ble of Contents Table of ContentsPreface Preface. . . . . . . . . . . . . .Systems Issues in Solid Freeform Fabrication SystemsIssuesin Solid Freeform FabricationMichael J. Wozny Michael J. Wozny Rensselaer Design Research Center, Rensselaer Polytechnic Institute Rensselaer Design Research Center, Rensselaer Polytechnic InstituteProcess Control System for a High Temperature. Process Control System for a High Temperature Workstation Performing the Selective Laser Sintering Process Workstation Performing the Selective Laser Sintering Process Suman Das, Richard H. Crawford, and J.J. Beaman, Suman Das, Richard H. Crawford, and J .J . Beaman, Dept. ofMechanical Engineering, The University ofTexas at Austin Dept. of Mechanical Engineering, The University of Texas at AustinA Parallel Slicing Algorithm for Solid Freeform A Parallel Slicing Algorithm for Solid Freeform Fabrication Processes Fabrication ProcessesC.F. Kirschman and C.C. Jara-Almonte, Centerfor Advanced C.F.

Beaman, Xiaoming (Michael) Deng, Guisheng Zong, and Joseph J. Beaman, Center for MS&E and Dept. ofMechanical Eng., The University ofTexas at Austin Center for MS&E and Dept. of Mechanical Eng, The University of Texas at Austin110 Design of a High Temperature Process Chamber for the Selective Design of a High Temperature Process Chamber for the Selective ...................... 110 Laser Sintering Process Laser Sintering ProcessJohn McWilliams, Christopher Hysinger, and J.J. Beaman, John McWilliams, Christopher Hysinger, and J.J. Beaman, Department ofMechanical Engineering, The University ofTexas at Austin Department of Mechanical Engineering, The University of Texas at AustinThe Application of an Artificial Body Force to the Selective Laser The Application of an Artificial Body Force to the Selective Laser Sintering Process Sintering Process118 ...................... 118Lawrence S. Melvin III and J.J. Beaman, Department ofMechanical Lawrence S. Melvin III and J .J.

118Lawrence S. Melvin III and J.J. Beaman, Department ofMechanical Lawrence S. Melvin III and J .J . Beaman, Department ofMechanical Engineering, The University ofTexas at Austin Engineering, The University of Texas at AustinCeramic Structures by Selective Laser Sintering of Micro.. ••••••.••••••••...•••••••••......•• 124 - 124 Ceramic Structures by Selective Laser Sintering of Micro- encapsulated, Finely·Divided Ceramic Materials encapsulated, Finely Divided Ceramic MaterialsN.K. Vail and J.W. Barlow, Department ofChemical Engineering N. K. Vail and J.W. Barlow, Department of Chemical Engineering The University ofTexas at Austin The University of Texas at AustinThe Measurement of the Thermal Properties and Absorptances The Measurement of Powders Near Their Melting Temperatures of Powders Near Their Melting Temperaturesof the Thermal Properties and Absorptances......... . .......0 0 0 0 0 0 0 0131 131Samuel S. Sih and Joel W.

. .......0 0 0 0 0 0 0 0131 131Samuel S. Sih and Joel W. Barlow, Department ofChemical Engineering Samuel S. Sih and Joel W. Barlow, Department of Chemical Engineering The University ofTexas at Austin The University of Texas at AustinMetal Parts from Selective Laser Sintering of Metal-Polymer Metal Parts from Selective Laser Sintering of Metal-Polymer Powders Powders141 - 141B. Badrinarayan and J.W. Barlow, Department ofChemical Engineering B. Badrinarayan and J.W. Barlow, Department of Chemical Engineering The University ofTexas at Austin The University of Texas at AustinDevelopment of a Selective Laser Reaction Sintering Workstation Development of a Selective Laser Reaction Sintering Workstation ............... B.R. Birmingham, J.V. Tompkins, G. Zong, andH.L. Marcus, B.R. Birmingham, J.V. Tompkins, G. long, and BL.

B.R. Birmingham, J.V. Tompkins, G. Zong, andH.L. Marcus, B.R. Birmingham, J.V. Tompkins, G. long, and BL. Marcus, Center for Materials Science and Engineering, The University ofTexas at Austin Center for Materials Science and Engineering, The University of Texas at Austin0 0 0 0 0 0 0 0147 147Application of Factorial Design in Selective Laser Sintering Application of Factorial Design in Selective Laser Sintering Xiaoming (Michael) Deng and Joseph J. Beaman, Xiaoming (Michael) Deng and Joseph J. Beaman, Department ofMechanical Engineering, The University ofTexas at Austin Department of Mechanical Engineering, The University of Texas at Austine e - - = =154 154Preliminary Experience with Selective Laser Sintigraphic (SLS) Preliminary Experience with Selective Laser Sintigraphic (SLS) Models of the Human Temporal Bone Models of the Human Temporal Bone.................0 0 0 0 0 0 0 0161 161Richard A. Levy, The University ofMichigan Medical Center and Richard A.

228 Have Different Scanner Geometries and Laser Spot Sizes Have Different Scanner Geometries and Laser Spot SizeslC. Nelson and J.W. Barlow, Department o/Chemical Engineering, J.C. Nelson and J.W. Barlow, Department of Chemical Engineering, The University o/Texas at Austin The University of Texas at AustinRapid Prototyping at Zero Gravity for In-Flight Repairs and Rapid Prototyping Fabrication on Space Station Freedom Fabrication on Space Station Freedom237 at Zero Gravity for In-Flight Repairs and ............... . .............. 237Floyd Roberts, Marshall Space Flight Center, Floyd Roberts, Marshall Space Flight Center, David Lomshek, Pittsburgh State University, and David Lomshek, Pittsburgh State University, and William E. Brower, Jr., Marquette University William E. Brewer, J12, Marquette UniversityMetal Parts Generation by Three-Dimensional Printing........................................

193Mukesh Agarwala, D. L. Bourell, B. Wu, and J. J. Beaman; Mukesh Agarwala, D . L . Bourell, B . Wu, and J . J . Beaman; The University of Texas at Austin The University of Texas at AustinSilicon Carbide Preforms for Metal Infiltration by Selective Laser Sintering™ Silicon Carbide Preforms for Metal Infiltration by Selective Laser SinteringTM 204 of Polymer Encapsulated Powders of Polymer Encapsulated Powders .......................................................................................... 204N. K. Vail, J. W. Barlow,and H. L. Marcus; N . K . Vail, J . W . Barlow,and H . L . Marcus; The University of Texas at Austin The University of Texas a t AustinFabrication of Ceramic and Metal Matrix Composites From Selective Laser From Selective Laser Fabrication of Ceramic and Metal Matrix Composites Sintered Ceramic Preforms Sintered Ceramic Preforms......................................................................................................

Representations Geometry Representations Crawford, and J. J. Beaman; Sashidhar Guduri, R. Sashidhar Guduri, R. H . Crawford, and J. J. Beaman; The University of Texas at Austin The University of Texas at Austin291 ....................................................................................................... 291Indirect Metal Composite Part Manufacture Using the SLS Process Indirect Metal Composite Part Manufacture Using the SLS Process .................................. 303 303 James R. Tobin, B. Badrinarayan, J. W. Barlow, J. J. Beaman, and James R. Tobin, B. Badrinarayan, J. W. Barlow, J. J. Beaman, and Bourell; The University of Texas at Austin D. D. L. Bourell; The University of Texas at AustinSolid Freeform.Fabrication ofSilicon Carbide Shapes by Selective Solid Freeform Fabrication of Silicon Carbide Shapes by Selective Laser Rea(!tion. Sint~ring.

325 325James V. Tompkins, and H. Marcus; James V . Tompkins, and H . L. Marcus; The University of Texas at Austin The University of Texas atAustinDrying ofColloidalllinder Infiltrated Ceramic Green Parts Produced Drying of Colloidal Binder Infiltrated Ceramic Green Parts Produced by Selective Laser SinteringTM ................................................................................................. 333 by Selective Laser Sintering™ 333M. Glazer, N.K. Vail, and J. W. Barlow; M . Glazer, N . K. Vail, and J. W . Barlow; The University. of TexasatAustin The University of Texas at AustinSynthesis~Selective Laser Sintering and Infiltration of HighTc DuaLPhase Synthesis, Selective Laser Sintering and Inﬁltration of High Tc Dual Phase Ag..YBa2Cu30,-x SupercondnctorComposites Ag-YBagCu307-X Superconductor339 ..................................................................... 339Composites Mukesh Agarwala, D. L. Bourell, A. Manthiram, B. R.

. . . . . . . . . . . . .311 - - -.-..311Thissell,W. Richards and Marcus, Harris L., The University of Thissell, W . Richards and Marcus, Harris L. , T h e University of Texas at Austin Texas at AustinMeasurementalld.Predictionof.the Thermal Conductivity Measurement and Prediction of the Thermal Conductivity of Powders.at High Temperatures of Powders at High Temperatures...................................................................................321 321Sih, SamuelS. and Barlow, Joel W., The University ofTexas at Sih, Samuel S. and Barlow, Joel W., The University of Texas at Austin AustinAnisotropy in Properties of Alumina Processed by SLS Anisotropy 1n Properties of Alumina Processed by SLS330 ................. 330Subramanian, P.Kamatchi, Vail, N.K., Barlow,J.W., and Marcus, Subramanian, P. Kamatchi, Vail, N. K., Barlow, J . W.

379 Bunnell, David, Bourell, David, and Marcus, Harris, The Bunnell, David, Bourell, David, and Marcus, Harris, The University of Texas at Austin University of Texas at Austin387 Modeling of Polymer Degradation in SLS Modeling of Polymer Degradation in SLS ............................................................................. 387Vail, N. K., and Barlow, J.W., The University of Texas at Austin Vail, N . K . , and Barlow, J . W . , The University of Texas at AustinUltraviolet Curing of Highly Loaded Ceramic Suspensions Ultraviolet Curing of Highly Loaded Ceramic Suspensions 396 For Stereolithography of Ceramics For Stereolithography of Ceramics ........................................................................................ 396Griffith, Michelle L. and Halloran, John W., The University of Griffith, Michelle L . and Halloran, John W.

Wozny Alejandro Beascoechea, Michael J. Wozny Rensselaer Polytechnic Institute Rensselaer Polytechnic InstituteIntegration of a Solid Freeform Fabrication Process into a Feature-Based CAD Integration of a Solid Freeform Fabrication Process into a Feature-Based CAD System Environment System EnvironmentYoung Suh, Michael J. Wozny Young Suh, Michael J. Wozny Rensselaer Polytechnic Institute Rensselaer Polytechnic InstituteFree Form Fabrication of High Strength Metal Components and Dies Free Form Fabrication of High Strength Metal Components and Dies+ C . C . Bampton, *R. Burkett +C.C. Bampton, *R. Burkett +Rockwell Science Center, *Rocketdyne +Rockwell Science Center, *RocketdyneEffect of Particle Size on SLS and Post-Processing of Alumina with Polymer Effect of Particle Size on SLS and Post-Processing of Alumina with Polymer Binders B i n d e r s+ P . Kamatchi Subramanian, + J . W . Barlowf H.L. Marcus +P. Kamatchi Subramanian, +J.W. Barlow,* H.L.

Marcus +P. Kamatchi Subramanian, +J.W. Barlow,* H.L. Marcus +The University of Texas at Austin, *University of Connecticutt “The University of Texas at Austin, *University of ConnecticuttModeling of Dynamic Effects Caused by the Beam Delivery System in Selective Modeling of Dynamic Effects Caused by the Beam Delivery System in Selective Laser Sintering Laser SinteringIrem Turner, Kristin Wood, Ilene Busch-Vishniac Irem Turner, Kristin Wood, Ilene Busch-Vishniac The University of Texas at Austin The University of Texas at AustinOptimizing Part Quality with Orientation Optimizing Part Quality with Orientation David Thompson, Richard Crawford David Thompson, Richard Crawford The University of Texas at Austin The University of Texas at AustinSelective Laser Sintering Preparation and Tribological Testing of Nanostructured Selective Laser Sintering Preparation and Tribolo‘gical Testing of Nanostructured Tungsten Carbide-Cobalt Composites Tungsten Carbide-Cobalt Composites+Rand D.

EFACEThe Eleventh Solid Freeform Fabrication (SFF) Symposium, held at The University of Texas in Austin on August 8-10, 2000, was attended by over 100 national and international researchers. Papers addressed SFF issues in computer software, machine design, materials synthesis and processing, and integrated manufacturing. New sessions on ceramic materials and multiple materials SFF were added to this year’s program. The diverse domestic and foreign attendees included industrial users, SFF machine manufacturers, university researchers and representatives from the government. The Symposium organizers look forward to its being a continuing forum for technical exchange among the expanding body of researchers involved in SFF.The Symposium was again organized in a manner to allow the multi-disciplinary nature of the SFF research to be presented coherently, with various sessions emphasizing computer issues, machine topics, and the variety of materials aspects of SFF.

We believe that documenting the changing state of SFF art as represented by these Proceedings will serve both the people presently involved in this fruitful technical area as well as new researchers and users entering the field.This is our first attempt at a CD proceedings of the meeting. We have provided hard copies of the front and back matter. Linked versions are included on the CD. The papers are stored individually on the CD in pdf format by primary author last name, and Adobe® Acrobat® Reader™ installers for the Macintosh and PC (Windows 2000) are included which may be used to view and search the pdf files. The Table of Contents file has links to all the papers. As shown on the Table of Contents, we have sequentially numbered the pages of the papers to facilitate citation.Adobe ® Acrobat ® Reader ™ has document search capabilities. To do a search, select “Search>Query” from the Edit menu.

To do a search, select “Search>Query” from the Edit menu. Prior to searching, add the CD index “index.pdx” (Mac) or “index” (PC) by clicking the Indexes button on the Query window to activate the SFF Proceedings index. Some versions of Reader™ do not have search capabilities which are necessary to keyword search the SFF Symposium Proceedings. If you have problems with searching, you might consider installing the version of Reader™ from the CD. For non-Macintosh/non-Windows 2000 platforms, you can download Adobe® Acrobat® Reader™ from the Adobe website (http://www.adobe.com/) free of charge; be sure to check the box indicating search capabilities before downloading.The editors would like to extend a warm “Thank You” to Rosalie Foster for her detailed handling of the logistics of the meeting and the Proceedings, as well as her excellent performance as registrar and problem solver during the meeting.

We would like to thank the Organizing Committee, the session chairs, the attendees for their enthusiastic contributions, and the speakers both for their significant contribution to the meeting and for the relatively prompt delivery of the manuscripts comprising this volume. We look forward to the continued close cooperation of the SFF community in organizing the Symposium. We also want to thank the Office of Naval Research (N00014-00-1-0674) and the National Science Foundation (DMI-0003777) for supporting this meeting financially. The meeting was co-organized by the University of Connecticut at Storrs, and the Mechanical Engineering Department, Laboratory for Freeform Fabrication and the Texas Materials Institute at The University of Texas at Austin.The editors.vii.

ble of ContentsPage # Preface ‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐ xOrganizing/Advisory Committee ‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐ xiSocietal ImpactAdditive Manufacturing Education in the UKPhill Dickensa, Phil Reevesb, Richard Haguec, aLoughborough University, bEconolyst Ltd., cUniversity of Nottingham ‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐ 1Innovative Education in Additive Manufacturing in ChinaFeng Lin, Lei Zhang, Ting Zhang, Jia Wang, Renji Zhang, Tsinghua University ‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐ 14Interactions of an Additive Manufacturing Program with SocietyFrank Liou, Ming Leu, Robert Landers, Missouri University of Science and Technology ‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐ 45GT MENTOR: A High School Education Program in Systems Engineeringand Additive Manufacturing David W.

Leu, Missouri University of Science and Technology ‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐ 734viTheoretical Modeling and Experimental Characterization of StressDevelopment in Parts Manufactured Through Large Area Maskless Photopolymerization Tao Wu, Suman Das, Georgia Institute of Technology ‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐ 748Evaluations of Temperature Measurements by Near­Infrared Thermography in Powder­Based Electron­Beam Additive Manufacturing Steven Pricea, Ken Cooperb, Kevin Choua, aThe University of Alabama, bMarshall Space Flight Center ‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐ 761Numerical Thermal Analysis in Electron Beam Additive Manufacturingwith Preheating Effects Ninggang Shen, Kevin Chou, The University of Alabama ‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐‐ 774Computational Modeling and Experimental Validation ofMicrostructural Development in Superalloy CMSX­4 Processed Through Scanning Laser Epitaxy Ranadip Acharya, Rohan Bansal, Justin J.

Landers, and Edward C. KinzelEffects of Inter-Layer Time Interval on Temperature Gradients in Direct Laser Deposited Ti-6Al-4V ................................ 175Mohammad Masoomi, Scott M. Thompson, Nima Shamsaei, and Linkan BianPlanning the Process Parameters during Direct Metal Deposition of Functionally Graded Thin-Walled Parts Based on a 2D Model ................................................................................................................................................................................ 185Jingyuan Yan, Ilenia Battiato, and Georges M. FadelAn Investigation into Melt Pool Effective Thermal Conductivity for Thermal Modeling of Powder-Bed Electron Beam Additive Manufacturing ..........................................................................................................................................................

605Rakish Shrestha, Jutima Simsiriwong, Nima Shamsaei, Scott M. Thompson, and Linkan BianviiMicrostructural and Mechanical Performance Of Al₂O₃ Nanoparticle Reinforced 17-4 PH Stainless Steel Bulk Composite Parts Fabricated by Laser Engineered Net Shaping Process ................................................................................. 617Fuda Ning, Yingbin Hu, Zhichao Liu, Hui Wang, Weilong Cong, and Yuzhou LiInvestigation of Heat-Affected 304L SS Powder and Its Effect on Built Parts in Selective Laser Melting ........................... 625Caitlin S. Kriewall, Austin T. Sutton, Ming C. Leu, Joseph W. Newkirk, and Ben BrownElevated Temperature Microstructure Stability of SLM 304L Stainless Steel ....................................................................... 640Tarak Amin and Joseph NewkirkMulti-Objective Process Optimization of Additive Manufacturing: A Case Study on Geometry Accuracy Optimization .... 656Amir M.

Elliott, P. Nandwana, D. Siddel, and B.G. ComptonA Smooth Toolpath Generation Method for Laser Metal Deposition ................................................................................... 1038Renwei Liu, Zhiyuan Wang, Yunlu Zhang, Todd Sparks, and Frank LiouIntegration of Voxel Based and Source Based Representation for Computer Aided Design of Functional Gradient Materials ............................................................................................................................................................................... 1047Fangquan Wang, Frank Liou, and Todd SparksCharacterization of Feedstock in the Powder Bed Fusion Process: Sources of Variation in Particle Size Distribution and the Factors that Influence Them ..................................................................................................................................... 1057J. Whiting and J.

1057J. Whiting and J. FoxPreliminary Testing of Nanoparticle Effectiveness in Binder Jetting Applications .............................................................. 1069Alta C. Bailey, Abbey Merriman, Amelia Elliott, and Mufeed M. BastiContinuous Fiber Angle Topology Optimization for Polymer Fused Filament Fabrication ................................................. 1078R. Hoglund and D.E. SmithProcess DevelopmentOptimization of Low One-Photon Polymerization for Hydrostatic 3D Printing of Silicone Material .................................. 1093Dong Sung (Danny) Kim, Steven Thompson, Melissa Grunlan, and Bruce L. TaiLaser Line Scan Characterization of Geometric Profiles in Laser Metal Deposition ........................................................... 1103M. Gegel, A. Nisbett, D. Bristow, and R.G. LandersUnderstanding the Laser Sinterting of Polymers at Microscale Level by Using X-Ray Computed Tomography ................

1621Haijun Gong, Chong Teng, Kai Zeng, Deepankar Pal, Brent Stucker, J.J.S. Dilip, Jack Beuth, and John J. LewandowskiStatistical Process Control Application to Polymer Based SLS Process .............................................................................. 1634S.O. Akande, K.W. Dalgarno, J. Munguia, and J. PallariA Short Study on the Fabrication of Single Track Deposits in SLM and Characterization .................................................. 1644J.J.S. Dilip, Ashabul Anam, Deepankar Pal, and Brent StuckerVariation of Emittivity with Powder Bed Fusion Build Parameters ..................................................................................... 1660J.C. Heigel, B.M. Lane, and S.P. MoylanSpectral Analysis of Infrared Lamps for Use in the High Speed Sintering Process .............................................................

Friel, and Russell A. HarrisPerformance of Sulfur Concrete in Planetary Applications of Contour Crafting ................................................................. 2282Xiao Yuan, Jing Zhang, Behnam Zahiri, and Behrokh KhoshnevisxvImpact of Vapor Polishing on Surface Roughness and Mechanical Properties for 3D Printed ABS ................................... 2295Clayton Neff, Matthew Trapuzzano, and Nathan B. CraneSurface Scanning Methods and Large-Scale FDM Printng for the Replication of Watercraft Layup Tooling .................... 2305D.W. Nuttall, A.M. Elliott, B.K. Post, and L.J. LoveDevelopment of Novel Tapered Pin Fin Geometries for Additive Manufacturing of Compact Heat Exchangers ............... 2314Julien Cohen and David L. BourellEffect of Sparse-Build Internal Structure on Performance of Fused Deposition Modeling Tools under Pressure ............... 2337S. Meng, L. Mason, G. Taylor, X. Wang, M. C. Leu, and K.

115Brian Torries, Nima Shamsaei, and Scott ThompsonEffect of Specimen Surface Area Size on Fatigue Strength of Additively Manufactured Ti-6Al-4V Parts .......................... 122Jonathan Pegues, Michael Roach, R. Scott Williamson, and Nima ShamsaeiSmall-Scale Mechanical Properties of Additively Manufactured Ti-6Al-4V ...................................................................... 134M. Haghshenas, O. Totuk, M. Masoomi, S.M. Thompson, and N. ShamsaeiDesign and Fabrication of Functionally Graded Material from Ti to Γ-Tial by Laser Metal Deposition ............................. 148Xueyang Chen, Lei Yan, Joe Newkirk, and Frank LiouTailoring Commercially Pure Titanium Using Mo₂C during Selective Laser Melting ......................................................... 160X.P. Li, S. Dadbakhsh, K. Vanmeensel, J. Vleugels, J. Van Humbeeck, and J.P.

744Yuanjie Zhang, Bo Song, Lei Zhang, Zhiwei Wang, Yusheng ShiConstruction of Metallic Glass Structures by Laser-Foil-Printing Technology ................................................................... 755Yiyu Shen, Yingqi Li, and Hai-Lung TsaiBuilding Zr-Based Metallic Glass Part on Ti-6Al-4V Substrate by Laser-Foil-Printing Additive Manufacturing ............... 771Yingqi Li, Yiyu Shen, Ming C. Leu, and Hai-Lung TsaiOptimising Thermoplastic Polyurethane for Desktop Laser Sintering ................................................................................ 785I.P.F. Richards, T.M.N. Garabet, I.S. Bitar, and F.M. SalmonviiiModelingReal-Time Process Measurement and Feedback Control for Exposure Controlled Projection Lithography ......................... 801Xiayun Zhao and David W. RosenOptimization of Build Orientation for Minimum Thermal Distortion in DMLS Metallic Additive Manufacturing .............

1082Z. Wang and D.E. SmithSimulation of Planar Deposition Polymer Melt Flow and Fiber Orientaiton in Fused Filament Fabrication ...................... 1096B.P. Heller, D.E. Smith, and D.A. JackNumerical Investigation of Stiffness Properties of FDM Parts as a Function of Raster Orientation .................................. 1112Sanchita Sheth and Robert M. TaylorA Two-Dimensional Simulation of Grain Structure Growth within Substrate and Fusion Zone during Direct Metal Deposition ....................................................................................................................................................................... 1121Jingwei Zhang, Wei Li, Frank Liou, and Joseph NewkirkNumerical Simulation of Temperature Fields in Powder Bed Fusion Process by Using Hybrid Heat Source Model .........

1918Marlon Wesley Machado CunicoxiiiFabrication and Characterization of Ti6Al4V by Selective Electron Beam and Laser Hybrid Melting .............................. 1924Bin Zhou, Jun Zhou, Hongxin Li, and Feng LinDevelopment of a Hybrid Manufacturing Process for Precision Metal Parts ..................................................................... 1935Leon Hill, Todd Sparks, and Frank LiouDefects Classification of Laser Metal Deposition Using Acoustic Emission Sensor ......................................................... 1952Haythem Gaja and Frank LiouAn Online Surface Defects Detection System for AWAM Based on Deep Learning ........................................................ 1965Shangyong Tang, Guilan Wang, Haiou Zhang, and Rui WangDevelopment of Automatic Smoothing Station Based on Solvent Vapour Attack for Low Cost 3D Printers ....................

WickerDesign for Protection: Systematic Approach to Prevent Product Piracy during Product Development Using AM ............ 2481U. Jahnke, R. Koch, and A. T. OppermannThe Use of Electropolishing Surface Treatment on IN718 Parts Fabricated by Laser Powder Bed Fusion Process ........... 2493Li Yang, Chris O’Neil, and Yan WuTowards Defect Detection in Metal SLM Parts Using Modal Analysis “Fingerprinting” .................................................. 2503James Urban, Nick Capps, Brian West, Troy Hartwig, Ben Brown, Robert Landers, Douglas Bristow, and Edward KinzelElectrochemical Enhancement of the Surface Morphology and the Fatigue Performance of Ti-6Al-4V Parts Manufactured by Laser Beam Melting ............................................................................................................................. 2516S. Bagehorn, J. Wehr, S. Nixon, A. Balastrier, T. Mertens, and H. J.

2546Mohammad Masoomi, Scott M. Thompson, Nima Shamsaei, and Meysam HaghshenasHow Significant Is the Cost Impact of Part Consolidation within AM Adoption? ............................................................ 2551A. Stevenson, M. Baumers, J. Segal, and Sarah MacdonellMethod for the Evaluation of Economic Efficiency of Additive and Conventional Manufacturing ................................... 2563C. Auth, A. Arndt, and R. AnderlxviIntegrating AM into Existing Companies - Selection of Existing Parts for Increase of Acceptance .................................. 2575A. Kruse, T. Reiher, and R. KochRamp-Up-Management in Additive Manufacturing – Technology Integration in Existing Business Processes ................. 2585J. Büsching and R. KochRational Decision-Making for the Beneficial Application of Additive Manufacturing .....................................................

2737Jared Allison, Conner Sharpe, Carolyn Conner Seepersad, and Steven KubiakGeometrical Accuracy of Holes and Cylinders Manufactured with Fused Deposition Modeling ....................................... 2757F. Knoop and V. SchoeppnerNew Filament Deposition Technique for High Strength, Ductile 3D Printed Parts ........................................................... 2777Nicholas Rodriguez and Richard CrawfordApplied Solvent-Based Slurry Stereolithography Process to Fabricate High-Performance Ceramic Earrings with Exquisite Details .............................................................................................................................................................

EFACEThe Thirty-fourth Annual International Solid Freeform Fabrication (SFF) Symposium – An Additive Manufacturing Conference, was held from August 14-16, 2023. There were 589 registrants from 9 countries. The total number of scheduled oral and poster presentations was 513. The meeting consisted of a Monday morning plenary, 56 parallel technical sessions and a poster session. The plenary session included an excellent mixture of presentations from Industry, US National Laboratories, and Universities. The Plenary Session also included presentations from the recipients of the 2023 International Outstanding Young Researcher in Freeform and Additive Manufacturing Award Dr. Niocholas Meisel, from Penn State University and Anthony Rollett from Carnegie Mellon University who won the International Freeform and Additive Manufacturing Excellence (FAME) Award.Two special events took place.

First was a panel discussion on AM Research – Reflections, Best Practices and Future Directions Lunch Panel. This Panel was organized by Dr. Guhaprasanna from Penn State. The panel consisted of experienced AM researchers from federal agencies, industry, laboratories, and academia. The Panel consisted of Dr. Khershed Cooper from the National Science Foundation, Dr. Andelle Kuszal from the Office of Naval Research, Dr. David Leigh, former CTO from 3D Systems, Dr. Amy Elliot from Oak Ridge National Laboratory, and Dr. Denis Cornier from Rochester Institute of Technology. The discussion focused on (1) a short overview of the panel members AM journey, (2) best practices in AM research for early-career AM researchers, (3) future directions in AM research for senior researchers. Second was a small networking forum on Blacks in Additive Manufacturing organized by Dr. Chukwuzubelu Ufodike from Texas A&M University.

Second was a small networking forum on Blacks in Additive Manufacturing organized by Dr. Chukwuzubelu Ufodike from Texas A&M University. The forum offered insightful discussions, networking opportunities, and the chance to explore career advancements and research collaborations. This event provided a valuable platform for diverse professionals, faculty, and students in Additive Manufacturing (AM). Methods to encourage under-represented students and faculty to attend the conference were discussed.There are over XX papers in this conference proceedings. Papers marked “REVIEWED” in the title area were peer reviewed by two external reviewers. The Table of Contents file and Author-Attendee file have links to all the papers. We have sequentially numbered the pages of the papers to facilitate citation.

The Table of Contents file and Author-Attendee file have links to all the papers. We have sequentially numbered the pages of the papers to facilitate citation. Manuscripts for this and all preceding SFF Symposia are available for free download at the conference website: https://www.sffsymposium.org/; select the “Proceedings Archive” pull-down menu item.The editors would like to thank the Organizing Committee, the session chairs, the attendees for their enthusiastic participation, and the speakers both for their significant contribution to the meeting and for the relatively prompt delivery of the manuscripts comprising this volume. We are grateful to TMS conference management staff for their significant contributions to the meeting planning and proceedings production, particularly Jeff Gnacinski, Tess De Jong, Kellye Parsson, Colleen Madore, and Trudi Dunlap. We look forward to the continued close cooperation of the additive manufacturing community in organizing the Symposium.

We look forward to the continued close cooperation of the additive manufacturing community in organizing the Symposium. We also want to thank the National Science Foundation (CMMI- 2226705) for supporting this meeting financially by providing 75 student registration fee waivers. Funding from the Office of Naval Research is also gratefully acknowledged. The meeting was organized within the Walker Department of Mechanical Engineering and the Center for Additive Manufacturing and Design Innovation (CAMDI) at The University of Texas at Austin.Joe Beaman took over as the Chair of the Organizing Committee The next SFF Symposium is August 12-14, 2024 at the Hilton Austin Hotel in Austin, Texas USA. The conference website will become active in mid-January 2023.The editors.xix.

velet Analysis of Solid Objects: Applications in Layered ManufacturingMark D. Van Roosendaal, Peter Chamberlain, CharlesThomasUniversity of UtahABSTRACTIn this paper, we introduce two-dimensional discrete wavelet basis functions and theirapplication in the analysis and modeling of surface topography in layered manufacturing objects. In previous work, a one dimensional wavelet transform technique was developed to generate variable thickness layers. [1] For vertical edge layers Haar wavelet decomposition is used the slicing direction but is not useful in the slicing plane. For frequency analysis within the slicing plane, biorthogonal wavelets provide the desired analysis ability. When analyzing layered manufacturing with ruled edges a true 2-D transform is appropriate.

For frequency analysis within the slicing plane, biorthogonal wavelets provide the desired analysis ability. When analyzing layered manufacturing with ruled edges a true 2-D transform is appropriate. Two-dimensional wavelet analysis simultaneously controls the layer thickness as well as the density of control points requiredthe surface definition of each layer edge.INTRODUCTION1) What is frequency analysis in layered manufacturing?It is possible to analyze a 3-D solid object in terms of its spatial frequency content. Using spatial frequency analysis, regions of high complexity appear as high frequencies and flat or gently curved regions appear as low frequencies. Wavelet analysis allows spatially resolved frequency analysis, where the location of the high or low frequencies is identified. By operating on the wavelet representation of an object with custom frequency filters, it is possible to simulate the results of various layered manufacturing processes.

a) Haar waveletb) biorthogonal wavelet4351000back together to fonn the original signal. The wavelets must satisfy certain criteria in order to be valid basis functions. Among these criteria, wavelets must be oscillatory and decay quickly to zero. [4] Two wavelets that have been found useful in layered manufacturing are the Haar wavelet [5] and the biorthogonal wavelet shown in Fig. 1.Discrete wavelet analysis (DWT, IDWT) assumes that a signal f(x) has been sampled at equally spaced intervals. The sequence length N of the signal being analyzed detennines how many wavelet levels can be represented. [6] When N=2n there are n + 1 wavelet levels. The levels represent different frequencies in the signal. Fig. 2 demonstrates frequency decomposition of a signal into wavelet levels. Level al0 is the lowest frequency level and al is the highest frequency level.PROPOSED PROCESS OVERVIEWThe proposedprocess of analyzing solid objects for frequency content is outlined in Fig. 4.

Changes in the part radius (r) can therefore be analyzed in two dimensions: vertical (z) and horizontal (8). Fig. 3 shows how the part is unwrapped. [1] The discrete wavelet transform is extended to two dimensions, which are decomposed and filtered simultaneously. Decisions regarding the type of wavelet basis function and the type of filter are made. After filtering, an inverse discrete wavelet transform is then used to reconstruct the object. From the resulting reconstruction, error is evaluated.INVERSE WAVELET RECONSTRUCTION OF SYNTHESIZED IMAGEt.........ERROR EVALUATIONtCUTTER PATH POINT REDUCTION AND GENERATIONIFigure 4. An overview of the 2-D wavelet analysis procedure.FILTERINGTwo types of filters, low pass and threshold, have been examined for application in layered manufacturing.Low Pass FilterHere, all wavelet coefficients above a specified frequency level are removed.

After inverse transforming, the model contains no frequency components above the cutoff. The resulting model looks as if it were produced by a constant layer thickness LM technique. Fig. 5b. shows these results.Threshold FilterHere, all wavelet coefficients below a specified threshold value are removed. The resulting model looks as if it were constructed from variable thickness layers. Fig. 5c. shows this result. Because of the ability to independently vary the filtering of each frequency level, remove wavelet coefficients above a specified frequency level, or perform global filtering of all frequency levels, thousands of possible filter combinations can be constructed.437. "DfltlIpolntaO·~F0.1O.O~F -0.052004006008001000a.o"--_..~ o~~~..Data polm.Figure 5. a) This figure shows the Haar wavelet frequency coefficients for a single slice of the head. All ten wavelet levels are shown. Level dl is the highest frequency.

a) This figure shows the Haar wavelet frequency coefficients for a single slice of the head. All ten wavelet levels are shown. Level dl is the highest frequency. b) Results from low pass filtering, yielding uniform slice thickness. c) Results from threshold filtering, yielding adaptive slice thickness. Filtering was intentionally set high to better illustrate the results.POINT REDUCTIONWhen the original model is put intocylindrical format, the file consists of regularly spaced points as shown in Fig. 6a. Threshold filtering forces the points into straight lines, but the total number of points in the file remains constant. Deleting redundant points along each line segment reduces both point density and file SIze.,......a.. " " b.c.Figure 6. a) Point density prior to wavelet filtering. b) After wavelet filtering.

" " b.c.Figure 6. a) Point density prior to wavelet filtering. b) After wavelet filtering. c) After point reduction.438b.c.VERTICAL DIRECTION (z) RESULTSThe ability to analyze the object in the vertical direction ena~les the ?~erat~r to ~enerate variable thickness layers. In Figures 7 and 8, below, a single contour m the shcmg dIrection was analyzed.110.5109.5rr'r7rr'7'"6.5r-7"'7'"~7.56.56 500550600650700750a.b.Figure 7. a) Vertical slice (side view of front portion of the head) after wavelet filtering. Filtering was intentionally set high to better illustrate the results. b) Close up view of the nose. Notice the variable layer thickness. Gray represents the original contour; black represents the contour after processing. The Haar WT is ideal for vertical edge layered manufacturing. In this example the head was reconstructed from 76 variable thickness layers.11.

The Haar WT is ideal for vertical edge layered manufacturing. In this example the head was reconstructed from 76 variable thickness layers.11 . - - - - - - r - - - - , - - - , - - - - - - , - - - - ,10.5107.5~00;;:---5;;;50:----=600;:----;6=5O---=7oo::----c:!750a.b.Figure 8. a) Vertical slice (side view of front portion of the head) after wavelet filtering. Filtering was intentionally set high to better illustrate the results. b) Close up view of the nose. Notice the variable layer thickness. Gray represents the original contour; black represents the contour after processing. The biorthogonal WT is ideal for ruled edge layered manufacturing.

Notice the variable layer thickness. Gray represents the original contour; black represents the contour after processing. The biorthogonal WT is ideal for ruled edge layered manufacturing. In this example the head was reconstructed from 40 variable thickness layers.439HORIZONTAL (e) DIRECTION RESULTSThe ability to analyze the object in the e direction enables the operator to filter unwanted information (high frequency steps) in the curter path.02244668 -10864206108 -108642a.b.Figure 9. Original slice, gray (1024 points), slice after processing using filtered Haar wavelets and point reduction scheme, black. a) 784 points. b) 1020 points.2244668 -108642108 -1086420a.b.Figure 10. Original slice, gray (1024 points), slice after processing using threshold biorthogonal wavelets and point reduction scheme, black. a) 560 points. b) 630 points.The above figures (Fig.

Original slice, gray (1024 points), slice after processing using threshold biorthogonal wavelets and point reduction scheme, black. a) 560 points. b) 630 points.The above figures (Fig. 9,10) show how wavelet filtering provides a reduction in thenumber of control points defining the layer edge while preserving an accurate edge44061010representation. Figures 9 and 10 demonstrate the advanta~e of using th~ biorthogonal wavelet over the Haar the theta direction. Using Haar, 1020 pOints were requlfed to reproduce the original slice geometry while only 630 points were required using the biorthogonal wavelet. Based on the results shown, we anticipate this analysis win simplify and smooth the cutter path definitiona.b.c.Figure 11. a) Original resolution in the stacking direction).object. b) Predicted head aftertheta direction).

a) Original resolution in the stacking direction).object. b) Predicted head aftertheta direction). c) Predicted head afterWT filtering and IDWT (low WT filtering and IDWT (low resolution11 shows the original head and the reconstructed head after unwrapping, biorthogonal wavelet transform decomposition, threshold filtering, inverse transforming, and re wrapping. Fig. 11 (a) is the original STL file. It is assumed that the apparent layering effect is an artifact produced when the head was scanned. In Fig. 11 (b) the filters were set for fme resolution the layering direction (z) and low resolution around the circumference direction (8). Note that there is no apparent change in resolution in the z direction. The nose is broadened and the defmition around the ears is poor due to the more aggressive filter in the 8 direction. In Fig. 11 (c) the filters were set for low resolution in the layering direction (z) and fine resolution around the circumference direction (8).

In Fig. 11 (c) the filters were set for low resolution in the layering direction (z) and fine resolution around the circumference direction (8). Now thicker layers are apparent in the z direction. As expected, the low resolution filtering greatly affected the accuracy of the part causing the lips to be completely removed.LIMITATIONSare two primary limitations for the process outlined in this paper. 1) Currently, thistechnique works only for single-value functions. Decomposing complex geometry into simple sub-parts prior to processing can eliminate this limitation. 2) The sequence length N is chosen so that f(x) is a desirable representation of the original data. For simple, low-resolution parts, a lower value ofN (say, 128) is adequate. For higher resolution parts, a larger N (say, 1024) is recommended.

For simple, low-resolution parts, a lower value ofN (say, 128) is adequate. For higher resolution parts, a larger N (say, 1024) is recommended. Although a larger value ofN gives a better approximation, the value of this integer is limited by the physical constraints of the decomposition software or by the desire to perform computations effectively and efficiently. [5]441CONCLUSIONSThe 2-D wavelet transform is a useful tool for mathematical analysis of the spatialfrequency content of three-dimensional objects. This analysis enables the prediction appropriate layer thickness as well as reducing the cutter path point load for layered manufacturing processes. In order to perform wavelet analysis, 3-D objects are decomposed (unwrapped) into a cylindrical coordinate representation. This single valued representation is then wavelet transformed, filtered, and inverse transformed. The choice of filter (low-pass or threshold) is motivated by the application.

This single valued representation is then wavelet transformed, filtered, and inverse transformed. The choice of filter (low-pass or threshold) is motivated by the application. Uniform layer thickness part decomposition is created using low-pass filtering. Threshold filtering produces adaptively sliced representations.As expected, there is a direct relationship between filtering and error. Under-samplingand over-filtering were the most common sources of error. The Haar wavelet is useful for analyzing vertical edge layered manufacturing while the biorthogonal wavelet proved useful for may be ruled edges. addition, the useful for file compression.wavelet representation of the part is quite compactFUTURE WORKSeveral areas have been identified for further work. These areas include techniques forerror quantification and filter optimization. These techniques will allow user input of maximum error desired in both directions and will adjust the frequency filtering accordingly.

These techniques will allow user input of maximum error desired in both directions and will adjust the frequency filtering accordingly. Future work includes the development of more sophisticated software tools for analyzing, decomposing, and filtering stereolithography files and the creation of cutter paths. Future software will automate the process outlinedthis paper.REFERENCES1. Manufacturing." Ph.D. dissertation, University of Utah Dept. of Mechanical Engineering, 1997. Lee, C. H., Thomas, C.L., "Wavelet transform based analysis for layered manufacturing", Proceedings ofthe Seventh international Conference on Rapid Prototyping, 3/31-4/3, 1997. 3.Akansu, A. N., Smith, MJ.T, Subband and Wavelet Transforms, Design and Applications. Kluwer Academic Publishers, 1996, 55-82. 4. Young, R. K., Wavelet Theory and its Applications. Kluwer Academic Publishers, 1993, 1 5.

The system comprises a 1.2 kW C021aser, positioning system, beam scanner, non-contacting positioning sensor, beam conditioner and CAD/CAM system. This paper reports progress in relating machine parameters (scan rate, feed, beam power and polarization) to process measurables (material removal rate and surface roughness), and demonstrates the potential for rapid prototyping and direct manufacturing of: (a) rotationally symmetric components based on ablative ceramics such as Si3N4 and (b) graphite fuel cell plenums.A versatile, automated,INTRODUCTIONAdvanced ceramic and ceramic composite materials based on, for example, Si3N4 , SiC, SiAION, AIN, are candidate materials for automotive engine components, bearings, cutting tools, pump components, valves and valve-train components for conventional piston engines, 5 Currently, the high costs and turbochargers, recuperators, heat exchangers and hot gas filters.

I difficulties of shaping such brittle ceramics by diamond grinding, waterjet cutting, ultrasonic machining, sintering, and near-net shape processes, such as hot isostatic pressing, are major factors limiting their widespread deployment. In this research, the laser based system under development offers the potential not only for rapid prototyping but also for economical, direct manufacturing of full strength, structural ceramic components.BACKGROUNDin this program is based on patents developed by Copley, Bass and Hsu6 and by Bass and Copley7. More than a decade of research laser machining of complex, three-dimensional has shown that for optimized, reproducible, ceramic shapes, closed-loop control of beam mode, polarization, beam power, beam speed, beam feed, and laser firing in synchronization with translational, rotational and tilt stages will be required.

8-24 Discoveries by Copley, Bass and Hsu6 have identified conditions for removing layers of ablative ceramics with controlled wall angles (perpendicular or sloping) at all positions along the bounding wall of a closed contour. 6 Parameters for laser machining Si3N4 surfaces with arithmetic average surface roughnesses of 3 p.m,19 and post laser machining treatments, which restore the mechanical properties of the laser machined surfaces to those of the as-ground material; have been identified.

Note that,material surfaoeFigure 2: Beam scanner.Figure 3: Plan view for layer removal in a single quadrant to produce a piecewise linear approxima tion of a convex spherical contour. The beam is scanned in a single direction.306unlike solid free form fabrication, additional controls are incorporated in this system to select the terminal angle at which the laser beam contacts the component surface for wall slope control.

The system is capable of generating shapes by ablating material from oversized ceramic preforms or by solid free form fabrication for prototyping or direct manufacturing of components.MATERIALSThe materials used in the present study were Ceralloy 147-31E sintered reaction bonded Si3N4 supplied as blocks and 10 mm diameter bar, Allied Signal GS-44 Si3N4 (blocks), Noralide NCX-5102 HIPed Si3N4 (10 mm diameter bar), and nuclear grade graphite received from Oak Ridge National Laboratory.PARAMETRIC STUDIESMaterial removal rates for single grooves and layers machined in both Si3N4 and graphite were studied as a function of beam velocity, beam polarization, position of the focal plane and beam feed to groove width ratio (f/a). Figure 4 compares MRR data for Allied Signal GS-44,........, ...... I 00 C'J8 <.> C'J I 0 --t :xl ..........,24222018P == 940 W~0t::. A16Q) ....... cd ~ ........ <.d 12 P-o 8 10140©00Q) ~ ........ cd • 1""'4 ~ Q) .......

I 00 C'J8 <.> C'J I 0 --t :xl ..........,24222018P == 940 W~0t::. A16Q) ....... cd ~ ........ <.d 12 P-o 8 10140©00Q) ~ ........ cd • 1""'4 ~ Q) ....... cd :::g8640t::.25500GS-44,tBl 00---1:::.--- GS-44, tBl 90 0 NC-132, <3==0 0 NC 132, <3==900---j.---175 125 75 Scanning Veloci ty [em s-l]100150200225Figure 4: Comparison between material removal rates determined for Allied Signal Si3N4 #8, GS-44 MOR bars (present study) and Norton NC-132 MOR bars (Copley et al.) for beam velocities from 5 to 240 cm S·l, and electric vector parallel (0°) or perpendicular (90°) to the scan direction.307250Norton NC-132 collected by Wallace10 and Wallace etmachined at 940 W with prior data al,17 and shows that removal rates were approximately doubled for the GS-44 material. material removal rate of 15 x 10-3 cm seconds.

material removal rate of 15 x 10-3 cm seconds. Figures 5a, 5b show single layers machined in GS-44 with fla ratios of 0.9 and Twelve passes of the laser beam can be clearly seen fla ratio to 0.4 (Fig. 5b) resulted in a significantly smoother surface. Note that although entrance surfaces were covered with a layer of silicon in the above example, silicon build up was minimized during multi-layer removal through optimization of beam power, speed and feed.is equivalent to removing 1 cm3 of material(f/a :::: 0.9). ReductionFigure 5a: Single layer machined in GS-44 with fla ratio of 0.9.Figure 5b: Single layer machined in GS-44 with fla ratio of 0.3.Material removal of graphite was investigated at 940 W in the velocity range 10 tocm S-1 and· found to increase (a) with increasing velocity, and (b) as the position of the focal plane was changed from (Fig. 6).

6). Interestingly, the material removal rates were similar in magnitude to those for Si3N4 • groove and .layer morphologies were observed, material removal rates and surface roughnesses were optimized using a velocityto 1.0 or 1 mm below the specimen surfaceemplane position of -1 mm and fla of 0.35.308I fI.l Qj8()B7 c-P == 940WII...t"'J I 0 "I""""l ~ .........,Q) ~ c.O p::j ......-I c.O >- 0 86 -5 -... • AIIIQ) ~ ....... (.Q ....... H Q) ......,> (.Q ::g4 f-IIIIIIII ...

• AIIIQ) ~ ....... (.Q ....... H Q) ......,> (.Q ::g4 f-IIIIIIII ... •!:::,.Z == -0.5 mm !:::,.z == -1.0 mm 1.5mm 6.z ==51015 Scanning Velocity [cm s-l]20Figure 6: Material removal rates determined for graphite for beam scanning velocities from 5 to 20 cm s-t, electric vector parallel (0°) to the scan direction, and three different focal plane positions.APPLICATIONSRotationally Symmetric Si3N4 ComponentsThe material removal rate data collected above were applied to demonstrate feasibility of concept for laser machining rotationally symmetric Si3N4 components, such as valve stems and injector nozzles. Layers were laser machined from the 10 mm diameter Ceraloy 147-31E and Noralide NCX-5102 HIPed Si3N4 bars. It should be noted that the NCX-5102 HIPed bars were bowed, out of round and exhibited two seams and a flat on the external surface.

It should be noted that the NCX-5102 HIPed bars were bowed, out of round and exhibited two seams and a flat on the external surface. Figure 7a shows as-laser machined (940 W, 100 cm S-1) surfaces of NCX 5102, with the seam on the original surface labelled A. Regions B, and D were machined to cylinders, eliminating the surface features and out-of round shape. Note that region C was a shallower depth which was insufficient to eliminate the surface seam. Conditions which retained the surface profile with depth were also identified. Region E has been machined to the same depth as region D, but the surface seam is now clearly visible. Minimal silicon build was observed on the sample, unlike the continuous silicon films shown on individual layers of GS-44 in Figures 5a, 5b. Surfaces B, C and D also exhibited a smooth, brown film, thought to be a silicon oxy-nitride.

Surfaces B, C and D also exhibited a smooth, brown film, thought to be a silicon oxy-nitride. Residual silicon was removed by an etching treatment but the time was insufficient to completely remove the brown film (Figure 7b). Studies to optimize etching conditions are now in progress.30925Figure 7a: As-laser machined surfaces of Noralide NCX-5102 HIPed Si3N4•Figure 7b: Etched surfaces of Noralide NCX-5102 HIPed Si3N4•Fuel Cell PlenumsFuel cells convert chemical energy directly into electrical energy by oxidation of the fuel which flows through a complex array of channels in the plenum. Multiple plenums, typically made of graphite, are assembled to make the fuel cell stack. Each plenum contains channels approximately 1 mm in width, depth and spacing. Machining the complex array of channels is slow, produces carcinogenic graphite dust, and is expensive, even with CNC milling machines, since excessive force or speed results in chipping of the graphite.

Plenums containing developmental 1 x 1 inch arrays of channels were conventionally machined for lIT's fuel cell program at a cost of $500 each.In contrast, laser machining offers an economical and clean method for two- and three-dimensionally shaping graphite. Using the parameters developed above, a prototype fuel cell was machined as shown in Figure 8. Production costs, based on laser operating costs of $100.00 per hour, were estimated at approximately $8.00.SUMMARYThe above studies demonstrate the potential for direct manufacturing of laser machined components of Si3N4 and graphite. The research on fuel cells is being scaled up310Figure 8: Laser machining atechnology transfer to .11..11..11.,,""'"'''''''.11. production of complex""JlJl ....' ..... A.II.J."" system continuesREFERENCES1.Stinton,A.forced Composites by Chemical Energy Advanced Research Semiannual Progress TN, pp. 103-7. R. A.

S. M. Copley, M. Bass and R. Hsu: "Method and Apparatus for Shaping Articles Using a Laser Beam," U. S. Patent 4,914,270, April, 1990. M. Bass and S. M. Copley, "Noncontacting Position Sensor," U.S. Patent 4,888,490, December 19, 1989. S. M. Copley, M. Bass and R. J. Wallace: "Shaping Silicon Compound Ceramics with a Continuous Wave Carbon Dioxide Laser," Proc. Second Internat. Symposium on Ceramic Machining and Finishing," NBS, 1978, pp. 97-104. S. M. Copley, M. Bass and R. J. Wallace: "Shaping Silicon Compound Ceramics with a Continuous Wave Carbon Dioxide Laser," Proc. Second Internat. Symposium on Ceramic Machining and Finishing," NBS, 1978, pp. 97-104. R. J. Wallace: "A Study of the Shaping of Hot-Pressed Silicon Nitride with a High Power CW CO2 Laser," Ph. D. Dissertation, University of Southern California, Los Angeles, 1983. S. M. Copley, R. J. Wallace and M. Bass: "Laser Shaping of Materials," in Lasers in Materials Processing, A. Metzbower, ed. ASM Conf.

S. M. Copley, R. J. Wallace and M. Bass: "Laser Shaping of Materials," in Lasers in Materials Processing, A. Metzbower, ed. ASM Conf. Proc., 1983, pp.82-92. S. M. Copley, M. Bass and R. G. Wallace, "Shaping Silicon Compound Ceramics with a Continuous Wave Carbon Dioxide Laser," in The Science of Ceramic Machin and Surface Finishing II, eds. B. J. Hockey and R. W. Rice, NBS Special Publication 562, U. S. Government Printing Office, Washington, D.C., 1979, pp. 283 92. R. K. C. Hsu: "Shaping Articles by Vaporization with a CO2 Laser," Ph. D. Dissertation, Department of Mechanical Engineering, University of Southern California, Los Angeles, CA, April, 1984. S. M. Copley, "Laser Machining," in Encyclopedia of Materials Science, M. Bever,2511-12, Pergamon Press (1985).R. J. Wallace and S. M. Copley: Extended Abstracts - American Ceramic Society 87th Annual Meeting, Forum II - Ceramics for Heat Engines, Paper 8-FII-85, Cincinnati, Ohio, May, 1985. R. J. Wallace, M. Bass and S. M.

315-21.31220.21.22.23.24.H. Tao, W. Chang, S. M. Copley and J. A. Todd: "Effect of Surface Treat ment on the Strength of Laser Machined Silicon Nitride," in Laser Materials Processing III, J. Mazumder and K. N. Mukherjee, eds. The Minerals, Metals and Materials Society, AIME, 1989, pp. 199-215. R. K. C. Hsu and S. M. Copley: "Producing Three-Dimensional Shapes by Laser Milling," ASME Journal of Engineering for Industry, Vol. 112, No.4, 1990, pp. 375-79. S. M. Copley, HongYi Tao and J. A. Todd, "Post Treatment to Improve the Strength of Laser Shaped Silicon Compound Ceramics", U. S. Patent 5,022,957, June 11, 1991. J. A. Todd, S. M. Copley and R. C. Dix: "Laser Machining of Advanced Ceramic Materials," Invited Paper in "Beam Processing of Materials," eds. S. M. Copley and J. Singh, TMS Fall Meeting Symposium Proceedings, Chicago, II, October 1992, TMS, 1993, pp. 111-130. S. M.

S. M. Copley and J. Singh, TMS Fall Meeting Symposium Proceedings, Chicago, II, October 1992, TMS, 1993, pp. 111-130. S. M. Copley: Laser Processing Systems: An Overview: Invited Paper in "Beam Processing of Materials," eds. S. M. Copley and J. Singh, TMS Fall Meeting Symposium Proceedings, Chicago, II, October 1992, TMS, 1993, pp. 33-39.313.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference3-DIMENSIONAL CELLULAR AUTOMATA SIMULATION OF GRAIN STRUCTURE IN METAL ADDITIVE MANUFACTURING PROCESSESXuxiao Li and Wenda Tan Department of Mechanical Engineering, The University of Utah, Salt Lake City, UT 84112AbstractDistinct grain structures have been observed in Metal Additive Manufacturing (MAM) processes. These grain structures feature columnar grains which occasionally mix with equiaxed grains. The occurrence of these grain structures is not yet fully understood. In this work, direct laser deposition process is studied as a typical MAM process. A finite volume model is first implemented to obtain the thermal history. Next, the thermal history is fed into a Cellular Automata (CA) model to simulate the epitaxial and competitive growth through which the columnar grains are formed.

Next, the thermal history is fed into a Cellular Automata (CA) model to simulate the epitaxial and competitive growth through which the columnar grains are formed. Nucleation is included in the model to predict the generation of equiaxed grains, and is characterized by two nucleation parameters, the nucleation density and the critical undercooling. The simulation results show that both the nucleation parameters and process parameters can significantly affect the grain structure. The simulated grain structures examined on different planes can be significantly different, revealing the complexity of the 3-dimensional grain structures in MAM processes.1. IntroductionMetal additive manufacturing (MAM) systems are categorized by [1] into three general categories: the powder bed system, the powder feed system, and the wire feed system.

IntroductionMetal additive manufacturing (MAM) systems are categorized by [1] into three general categories: the powder bed system, the powder feed system, and the wire feed system. Despite a fair diversity in each of these systems, a typical MAM process can be described as follows: the metal material is constantly fed to a moving energy source (electron beam or laser); the energy source melts the metal, and the metal solidifies to become the built part once the energy source moves away. Then the process is repeated in a layer-by-layer fashion in which complex 3- Dimensional (3D) geometry can be achieved. So far MAM processes have been studied intensively with a wide variety of metal alloys such as stainless steel 316L, aluminum alloy AlSi10Mg, titanium alloy Ti6Al4V, and nickel alloy IN718.Certain distinct grain structure features have been observed in MAM which actually apply to different metal alloys and MAM systems.

Columnar grains (grains having large aspect ratio) are almost always observed. The size of columnar grains can depend on the power of the energy source [2]. The orientation along which columnar grains grow can be changed by adjusting the scanning direction of the energy source [3]. Equiaxed grains (grains having small aspect ratio) are also observed in MAM, usually at the surface of the built [4, 5], but can be dispersed among columnar grains [6], “sandwiched” between two layers of columnar grains [2], or located specifically on the border of the molten pool [7]. Strong grain texture is found from elongated columnar grains, while equiaxed grains tend to reduce grain texture [6-10]. Different grain structures can cause significant variance and anisotropy in the mechanical property of the built part [8].

Different grain structures can cause significant variance and anisotropy in the mechanical property of the built part [8]. Therefore, it is very important to understand the mechanism of the occurrence of these grain structures in order to control the built quality and to even tailor desired grain structures [10].1030In literature, there exist widely accepted explanations for the occurrence of columnar and equiaxed grains, such as that found in [11]. The columnar grains are considered to be caused by an epitaxial growth mechanism [12] in which all the existing grains keep growing along one “easy- growth” direction (<100> direction for body-centered-cubic and face-centered-cubic metals [12]). Since only one direction is favored for grain growth, grains can be “stretched” and become columnar. The occurrence of equiaxed grains is attributed to the heterogeneous nucleation ahead of the Solidification Front (SF) [13].

The occurrence of equiaxed grains is attributed to the heterogeneous nucleation ahead of the Solidification Front (SF) [13]. If the total volume of the equiaxed grains ahead of the SF is large enough, columnar grains can be blocked by the equiaxed grains (Columnar-to-Equiaxed Transition, or CET). Depending on the local temperature gradient, , the grain structure can be either columnar or equiaxed [14, 15]., and the SF velocity𝐺𝐺Simulation works [16-21] have been dedicated to quantifying the above mechanisms of grain structure evolution. Among them, Cellular Automata (CA) method has been successfully implemented to model the epitaxial growth of columnar grains. Ref. [16] reveals that the on the orientation toward which the columnar grain grows is complexly determined by the local SF. Ref. [17] shows a gradual enlargement of grain size in the building direction due to the competitive growth of the columnar grains.

Ref. [17] shows a gradual enlargement of grain size in the building direction due to the competitive growth of the columnar grains. By incorporating a classical nucleation model to CA method, the occurrence of equiaxed grains is also captured [17, 21]. The evolution of grain , structure is highly coupled with the temperature condition around the molten pool, also referred as the thermal history, which actually determines the local . However, and 𝑇𝑇(𝑥𝑥, 𝑦𝑦, 𝑧𝑧, 𝑡𝑡) modeling thermal history is a very challenging task due to the complication caused by 𝑉𝑉 the fluid flow in the molten pool [22] as well as the geometry of metal powders [23]. It is also found that the gas flow will influence the size of the molten pool [24].

It is also found that the gas flow will influence the size of the molten pool [24]. Therefore, different levels of simplifications are assumed in literature such as ignoring the fluid flow and treating the metal powder as a homogeneous continuum, as reviewed in [25].𝐺𝐺𝑇𝑇(𝑥𝑥, 𝑦𝑦, 𝑧𝑧, 𝑡𝑡)𝐺𝐺In this work, the laser direct deposition process is studied as a typical MAM process. A temperature simulation based on heat conduction is first implemented to obtain the thermal history. Then a 3D CA model combined with a nucleation model is used to capture the epitaxial growth of columnar grains as well as the occurrence of equiaxed grains. Compared to existing works, this work features a detailed discussion about the nucleation in MAM, as it turns out that nucleation conditions can significantly affect the grain structure. This work also reveals the complexity of the 3D grain structure in MAM while most of the existing work is limited to 2D.

This work also reveals the complexity of the 3D grain structure in MAM while most of the existing work is limited to 2D. Parametric studies of laser power and scanning speed are conducted to demonstrate their effects on grain structure.2. Model description2.1 Temperature modelingThe temperature modeling methodology is adapted from a finite volume model for laser direct deposition [26] where a Cartesian grid and a level-set formulation are used to capture the metal-gas interface. The level-set equation is written as (Equation 1):,𝜕𝜕𝜕𝜕 𝜕𝜕𝜕𝜕 + 𝐹𝐹𝑝𝑝|𝛻𝛻𝛻𝛻| = 01031𝑉𝑉(1)whereis the level-set function, defined as the “signed distance” to the metal-gas interface [27]. is defined positive on the gas region, negative on the metal region, and as zero on the metal-gas is the interface growth velocity due to powder addition,𝛻𝛻interface, as illustrated in Figure 1. 𝛻𝛻 which determines the evolution of(Figure 1).

𝛻𝛻 which determines the evolution of(Figure 1). 𝐹𝐹𝑝𝑝𝛻𝛻Figure 1: Illustration of the level-set functionand the interface growth velocity.𝑭𝑭𝒑𝒑𝝍𝝍 By ignoring fluid flow, the temperature is solved only from the heat conduction equation,which is formulated in a conservative form (Equation 2):,𝜕𝜕(𝜌𝜌𝜌𝜌)2𝑇𝑇 + 𝑆𝑆̇ is the specific internal energy,𝜕𝜕𝜕𝜕 = 𝑘𝑘𝛻𝛻where temperature, and 𝜌𝜌 , andis the is the source term of the energy conservation equation. The material properties, 𝑘𝑘 are defined according to the local phase (solid metal, liquid metal, mushy metal or gas). The material studied in this work is stainless steel (SS) 316L for the metal phase and argon for the 𝜌𝜌 is located on the metal- gas phase, the properties of which are given in Table 1.

The material studied in this work is stainless steel (SS) 316L for the metal phase and argon for the 𝜌𝜌 is located on the metal- gas phase, the properties of which are given in Table 1. The source term gas interface and is composed of several parts (Equation 3):is the density,is the thermal conductivity,𝑒𝑒,𝑆𝑆̇𝑒𝑒𝑘𝑘𝑆𝑆̇.𝑆𝑆̇ = 𝑆𝑆̇𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙 + 𝑆𝑆̇𝑐𝑐𝑐𝑐𝑐𝑐𝑐𝑐 + 𝑆𝑆̇𝑙𝑙𝑙𝑙𝑟𝑟 The four terms on the RHS of Equation 3 represent the energy source from the laser absorption as well as the energy loss from convection and radiation, respectively. is treated as a Gaussian distributed surface source term.

is treated as a Gaussian distributed surface source term. Further details about Equation 1-3 is referred to [26, 28].𝑆𝑆̇𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙Table 1: Material properties of stainless steel 316L and argon.PropertyStainless steel 316L ArgonSolidLiquidDensity (kg m-3) Specific heat (J kg-1 K-1) Thermal conductivity (W m-1 K-1)800050019.280005002091.7843120.44Liquidus temperature (K)1673Solidus temperature (K) Melting Latent heat (J kg-1)16482500001032𝑇𝑇(2)(3)The Cartesian grid used for the temperature modeling is shown in Figure 2. A non-uniform grid (covering both gas and metal phases) is applied. In this work, +Y is the vertical (building) direction and the laser scans in the XZ plane.

A non-uniform grid (covering both gas and metal phases) is applied. In this work, +Y is the vertical (building) direction and the laser scans in the XZ plane. The laser scans only in the fine mesh region (the zoom-in region in Figure 2, with the mesh size being 25 µm) at the center of the calculation domain to guarantee the accuracy of the thermal history near the molten pool.Figure 2: The Cartesian grid used for temperature simulations.2.2 Grain structure modelingThe thermal history obtained from temperature model is then used as the input for the grain structure simulations. A 3D CA model based on a decentered square method is used to model the grain growth. A detailed description of this model can be referred to [28-30]. A uniform Cartesian grid is used for the CA model, with a mesh size of 3 µm, which is about one tenth of the typical size of substrate grains (~30 µm). The substrate grains are equiaxed grains randomly generated before a simulation starts.

The substrate grains are equiaxed grains randomly generated before a simulation starts. The local growth velocity of each grain is assumed to be a function of the local undercooling, referred as the “grain growth kinetics” (Equation 4):,−5𝑉𝑉𝑔𝑔 = 𝑎𝑎(𝛥𝛥𝑇𝑇)3+ 𝑏𝑏(𝛥𝛥𝑇𝑇) −42+ 𝑐𝑐𝛥𝛥𝑇𝑇 + 𝑑𝑑−3−4𝑎𝑎 = 1.091 × 10, 𝑏𝑏 = −2.034 × 10 is the local growth velocity (unit:where, 𝑐𝑐 = 2.740 × 10, 𝑑𝑑 = 1.151 × 10);is the local undercooling defined asis the local temperature. The (unit: 𝑉𝑉𝑔𝑔 𝛥𝛥𝑇𝑇 = grain growth kinetics (Equation 4) is obtained by a polynomial interpolation from the results in 𝑇𝑇𝑚𝑚 − 𝑇𝑇 𝑇𝑇𝑚𝑚 [29], as is the common practice in CA method [30].), whereis the metal liquidus temperature and 𝑚𝑚/𝑠𝑠𝛥𝛥𝑇𝑇𝐾𝐾𝑇𝑇In the current model, there are two nucleation mechanisms based on which grains start their growth: the epitaxial nucleation on the SF and the heterogeneous nucleation ahead of the SF.

On the SF, nucleation takes place by attaching the liquid atoms onto the partially melted grains on the fusion line such that the attached atoms are arranged in the same crystal orientation as those of the partially melted grains. This nucleation mechanism is referred as the epitaxial nucleation and is assumed to occur immediately once the local temperature on the SF drops below the liquidus temperature. Epitaxial growth refers to the grain growth initiated by epitaxial nucleation. Ahead1033(4)of the SF, heterogeneous nucleation can take place at “substrates” such as the impurities in the liquid metal, where the activation energy of nucleation is reduced. The liquid atoms are then attached to the substrates and are assumed to be arranged in a random crystal orientation. This nucleation mechanism is referred as heterogeneous nucleation in this work.

The liquid atoms are then attached to the substrates and are assumed to be arranged in a random crystal orientation. This nucleation mechanism is referred as heterogeneous nucleation in this work. It is noted that in epitaxial nucleation the growing grains only “inherit” existing crystal orientations while in heterogeneous nucleation the growing grains have new orientations.In this work, heterogeneous nucleation is modeled using the continuous nucleation model [31]. In this model, the nucleation rate are continuous functions of and the number of nuclei the undercooling, as shown in Figure 3. At a critical undercooling the number of nuclei changes rapidly from zero to the maximum value . The nucleation rate exhibits a peak at a critical undercooling while maintains zero when the undercooling being too large or too small.

The nucleation rate exhibits a peak at a critical undercooling while maintains zero when the undercooling being too large or too small. The nucleation rate, as a function of undercooling, can be considered as a Gaussian distribution, characterized by the mean undercooling , which can be determined by experimental observations [32]. 𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐𝐼𝐼𝑁𝑁𝑁𝑁0, and the standard deviation𝛥𝛥𝑇𝑇𝑁𝑁,𝑙𝑙𝑠𝑠𝑔𝑔Figure 3: Illustration of continuous nucleation model.The method in [33] is used to implement the continuous nucleation model. The three illustrated in Figure 3 are first determined for the nucleation. parameters Then according to random cells in the CA model are selected as potential locations for nucleation to occur; for each potential nucleation location, a critical undercooling is assigned based 𝑁𝑁0 on the Gaussian distribution characterized by.

When a potential nucleation cell is still liquid and the local undercooling is larger than the assigned critical nucleation undercooling of the cell, this cell will change from liquid to a nucleus, from which a new grain starts growing. The grain growth after epitaxial nucleation and heterogeneous nucleation is treated in the same way based on the decenter square method and the growth kinetics in Equation 4.,, and𝑁𝑁0𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐𝛥𝛥𝑇𝑇𝑁𝑁,𝑙𝑙𝑠𝑠𝑔𝑔and𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐𝛥𝛥𝑇𝑇𝑁𝑁,𝑙𝑙𝑠𝑠𝑔𝑔3. Results and Discussion3.1 Thermal historyA typical single-pass simulation result (scanning direction +X) is shown in Figure 4. Figure 4a shows a 3D view of the zero-level-set surface flooded by the temperature. The zero-level-set surface is the metal-gas interface which indicates the geometry of the built track (length: 3.5mm).1034Figure 4b-d show a three-view drawing of the region near the molten pool to better demonstrate the molten pool geometry (marked by solid black line).

From the top view (Figure 4b) the molten pool is of the shape of a teardrop elongated along the X direction, as similar to those observed in welding processes [12]. Some geometric definitions of the molten pool referred in this work are given as follows: the point with the highest temperature on the zero-level-set isosurface is referred as the laser center (the orange circle), and the XY plane that contains the laser center is referred as the center plane (the red dashed line). The “tail” of the molten pool (the blue circle) is defined as the tip of the teardrop on the top view. The “bottom” of the molten pool (the blue triangle) is the deepest position of the molten pool. The “fusion surface”, marked by the yellow dashed line, is defined as the part of the molten pool surface that will instantly solidify (temperature dropping below the liquidus temperature) when laser further scans in the +X direction.Figure 4: Typical single-pass temperature simulation results.

(a) is the 3D view of zero-level-set isosurface (metal-gas interface) flooded with temperature distribution. (b)-(d) are the top view, side view and front view of the molten pool (boundary marked by black solid line). The orange circle marks the laser center; the red dashed line marks the center plane; the yellow dotted line marks the fusion surface; the blue circle marks the molten pool tail; the blue triangle marks the molten pool bottom.in Equation 3) and the powder addition ( in Equation 1) to different locations at different times according to the laser speed and scanning pattern. Shown in Figure 5 is an example in which the laser scans with a zigzag pattern along the X direction. This process can be repeated for multi- pass, multi-layer simulations.Multi-pass simulation is implemented by applying the energy source (and𝑆𝑆̇𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙𝐹𝐹𝑝𝑝1035𝑆𝑆̇𝑙𝑙𝑟𝑟𝑟𝑟Figure 5: 3D view of zero-level-set isosurface flooded with temperature distribution in multi-pass simulations.

The laser scans in a zigzag way along the X direction. (a)-(d) show the simulation results at the end of the second, third, fourth and fifth pass. The simulation result at the end of the first track is shown in Figure 4a.3.2 Effects of maximum nucleation density and mean critical undercoolingA series of single-pass simulations (scanning direction +X) are carried out to investigate . the effects of the maximum nucleation density Due to a lack of experimental calibration of have and 𝑁𝑁0 𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐 been tested and they are found to have a significant influence on the grain structure. The laser powerand the mean critical undercooling , a range ofand𝑁𝑁0𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐𝑁𝑁0𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐and scanning velocityused in this section are 200 W and 26 mm/s.𝑉𝑉𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙𝑃𝑃 In order to better rationalize the simulation results, Hunt’s analytical model [13] is used to compare with the simulation results.

In Hunt’s model, G-V maps are analytically computed to show the combinations that lead to grain structures of either completely columnar, completely equiaxed, or “mixed” with columnar and equiaxed grains, as shown in Figure 5a. The and m/s G-V maps discussed in this work are all in logarithm with unit conventions of K/m for for tend to cause columnar grains while smaller 𝑉𝑉(𝐺𝐺, 𝑉𝑉). It can be seen in Figure 6a that largerand smaller𝐺𝐺and largertend to cause equiaxed grains.𝐺𝐺𝑉𝑉𝐺𝐺𝑉𝑉1036of 1014 m-3 and combinations that lead to columnar, equiaxed and mixed 𝑵𝑵𝟎𝟎 (𝑮𝑮, 𝑽𝑽) on the fusion surface shown in (b) and (c). The orange grain structures. The blue dots are the calculated 𝜟𝜟𝑻𝑻𝑵𝑵,𝒎𝒎𝒎𝒎𝒎𝒎𝒎𝒎 at the tail polygon marks the region covered by the blue dots. The blue circle and triangle indicate the distributions on the fusion surface observed from and and bottom of the molten pool. (b) and (c) are the top view.

In Equation 5b, is the angle between the laser scanning direction (+X) and the local normal direction of the fusion surface. Figure 6b and 6c plot on the fusion surface from a top view. It is observed that along the the distributions of center plane (red dashed line), the smallest are found at the tail of the molten pool 𝐺𝐺 (blue circle), while the largest are found at the bottom (blue triangle). Assuming that the normal directions of the fusion surface represent the easy growth directions of grain, it is expected that the equiaxed grains tend to occur at the tail and columnar grains tend to dominate at the bottom. This trend is also demonstrated in Figure 6a as the blue circle (tail) already reaches the completely equiaxed region while the blue triangle (bottom) stays in the completely columnar region.𝑐𝑐andand largest𝑉𝑉and the smallest 𝐺𝐺𝑉𝑉𝐺𝐺𝑉𝑉is varied as 1013 m-3, 1014 m-3 and 1015 m-3 while is kept as 5 K.

In this work, the simulated grain structure on a certain plane is represented by inverse pole figure (IPF) orientation map in which the “projection direction” (the direction perpendicular to the plane under examination) is mapped to the crystal frames. The color key of the IPF mapping used in this work is shown in Figure 7. The simulated grain structure of the center plane from these single-pass simulations is shown in Figure 8a-c. One increases. There are almost no can observe that the number of equiaxed grains increases as equiaxed grains visible in Figure 8a; columnar grains grow epitaxially from the substrate (the blackNow the effects ofon the grain structure are investigated.𝑁𝑁0 𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐𝑁𝑁0𝑁𝑁01037(5a)(5b)(5c)dashed line) toward the laser scanning direction (SD).

Equiaxed grains are found interspersed among the dominating columnar grains when is as large as 1015 m-3 (Figure 8c), the equiaxed grains block the columnar grains which grow epitaxially from 𝑁𝑁0 the substrate; a CET can be clearly observed. Figure 8d shows the G-V map calculated based on Hunt’s model with different condition on the fusion surface is shown values, and the is 1013 m-3, the again with the orange polygon. It confirms the results in Figure 8a-c: when 𝑁𝑁0 is 1015 m-3, a fusion surface is in the mixed grain region and some is in the completely equiaxed significantly portion of region.is 1014 m-3 (Figure 8b). When𝑁𝑁0(𝐺𝐺, 𝑉𝑉)barely enters the mixed grain region; in contrast when 𝑁𝑁0𝑁𝑁0(𝐺𝐺, 𝑉𝑉)(𝐺𝐺, 𝑉𝑉)Figure 7: The color key for the inverse pole figure orientation mapping.Figure 8: The effects of center plane with values of 𝑵𝑵𝟎𝟎 of fusion line. (d) is the G-V map obtained with different values of on the fusion surface.on grain structure.is kept as 5 K.

(d) is the G-V map obtained with different values of on the fusion surface.on grain structure.is kept as 5 K. (a)-(c) are the grain structures on the being 1013 m-3, 1014 m-3 and 1015 m-3. The black dashed line indicates the location𝜟𝜟𝑻𝑻𝑵𝑵,𝒎𝒎𝒎𝒎𝒎𝒎𝒎𝒎; the orange polygon stands for the𝑵𝑵𝟎𝟎𝑵𝑵𝟎𝟎(𝑮𝑮, 𝑽𝑽)The effects ofis studied next.is varied as 0 K, 5 K, and 10 K whileis fixed as 1014 m-3. The simulated grain structure on the center plane and the G-V map is shown 𝑁𝑁0 𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐 𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐 in Figure 9. It is observed in Figure 9a-c that a larger causes less equiaxed grain and correspondingly more columnar grains. This point is confirmed in Figure 9d as a decreasing portion of the increases.𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐polygon will be covered by the mixed or equiaxed region as(𝐺𝐺, 𝑉𝑉)𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐1038is kept as 1014 m-3. (a)-(c) are the grain structures on being 0 K, 5 K and 10 K.

(a)-(c) are the grain structures on being 0 K, 5 K and 10 K. The black dashed line indicates the ; the orange polygon standsFigure 9: The effects of the center plane with the values of location of fusion line. (d) is the G-V map obtained with different values of for theon grain structure.𝜟𝜟𝑻𝑻𝑵𝑵,𝒎𝒎𝒎𝒎𝒎𝒎𝒎𝒎𝑵𝑵𝟎𝟎𝜟𝜟𝑻𝑻𝑵𝑵,𝒎𝒎𝒎𝒎𝒎𝒎𝒎𝒎on the fusion surface.𝜟𝜟𝑻𝑻𝑵𝑵,𝒎𝒎𝒎𝒎𝒎𝒎𝒎𝒎(𝑮𝑮, 𝑽𝑽) It is concluded through these parametric studies that both largeencourage the domination of columnar grain, while small occurrence of equiaxed grain.and 𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐and smallwill will favor the𝑁𝑁0𝑁𝑁03.3 Effects of laser power and scanning velocityThe effects of laser power pass simulations using the three sets of as 5 K. The simulated grain structure on the center plane and the G-V map is shown in Figure 10.

The simulated grain structure on the center plane and the G-V map is shown in Figure 10. 𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐and scanning velocityare studied by carrying out single- is fixed as 1014 m-3 andin Table 2.𝑉𝑉𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙 𝑁𝑁0𝑃𝑃(𝑃𝑃, 𝑉𝑉𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙)Table 2: Process parameters used in section 3.3SetLaser power (W)Scanning velocity (mm/s)12002624002632008.81039is kept as 1014 m-3 and andas 5 K. (a)-(c) are = 26 𝜟𝜟𝑻𝑻𝑵𝑵,𝒎𝒎𝒎𝒎𝒎𝒎𝒎𝒎 = 8.8 mm/s. The black dashed line indicates 𝑽𝑽𝒍𝒍𝒎𝒎𝒍𝒍𝒎𝒎𝒍𝒍 𝑷𝑷 𝑷𝑷 ; the three on the fusion surface computed with different 𝑵𝑵𝟎𝟎 = 𝟏𝟏𝟎𝟎on grain structure.andFigure 10: The effects of the grain structures on the center plane with different the values of 𝑽𝑽𝒍𝒍𝒎𝒎𝒍𝒍𝒎𝒎𝒍𝒍 = 26 mm/s; (c) for mm/s; (b) for the location of fusion line.

(d) is the G-V map obtained with 𝑽𝑽𝒍𝒍𝒎𝒎𝒍𝒍𝒎𝒎𝒍𝒍 polygons of orange, blue and green color stand for the values of: (a) for= 200 W,𝑵𝑵𝟎𝟎 = 200 W,𝑷𝑷= 400 W,𝑽𝑽𝒍𝒍𝒎𝒎𝒍𝒍𝒎𝒎𝒍𝒍 −𝟑𝟑and𝑽𝑽𝒍𝒍𝒎𝒎𝒍𝒍𝒎𝒎𝒍𝒍𝑷𝑷𝑷𝑷 and𝟏𝟏𝟏𝟏𝜟𝜟𝑻𝑻𝑵𝑵,𝒎𝒎𝒎𝒎𝒎𝒎𝒎𝒎 = 𝟓𝟓𝟓𝟓𝒎𝒎.(𝑮𝑮, 𝑽𝑽)𝑽𝑽𝒍𝒍𝒎𝒎𝒍𝒍𝒎𝒎𝒍𝒍𝑷𝑷 The effects ofis kept at 26mm/s while is varied from 200 W (Figure 10a) to 400 W (Figure 10b). It is observed that increasing the laser power will encourage the occurrence of equiaxed grain, as columnar grain still dominates in Figure 10a but a CET is clearly observed in Figure 10b. This can be explained by comparing the G-V polygons (blue and orange) of the two cases in Figure 10d. These two regions is applied in these two cases.

This can be explained by comparing the G-V polygons (blue and orange) of the two cases in Figure 10d. These two regions is applied in these two cases. However, are comparable in terms of the largest a higher power (blue polygon) gives a much smaller minimum on the fusion surface, which encourages the occurrence of equiaxed grain especially at the tail region; therefore, a CET is observed in Figure 10b.can be seen by comparing Figure 10a and 10b in which𝑉𝑉𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙𝑃𝑃𝑃𝑃, as the same𝑉𝑉𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙𝑉𝑉𝐺𝐺is fixed as 200 W while is varied from 26 mm/s (Figure 10a) to 8.8 mm/s (Figure 10c). Reducing the scanning velocity tends to eliminate the occurrence of equiaxed grains, as in Figure 10a interspersed equiaxed grains are still visible while in Figure 10c grains are completely columnar. Again, this point can be explained in Figure 10d by comparing the green and orange polygons. The two polygons cover a common range of than the orange one.

Again, this point can be explained in Figure 10d by comparing the green and orange polygons. The two polygons cover a common range of than the orange one . Therefore, the green polygon is entirely within the completely columnar due to smaller 𝐺𝐺 region, leading to the completely columnar grain structure in Figure 10c.The effects ofcan be seen by comparing Figure 10a and 10c, as𝑉𝑉𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙𝑃𝑃𝑉𝑉𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙, but the green polygon has a smaller𝑉𝑉𝑉𝑉𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙It has been demonstrated through this series of simulations that the process parameters can significantly change the grain structuring via altering the local condition.

It is foreseeable that an effective simulation can help make “a priori selection of process parameters for tailored microstructure [34]” and (partially) replace the trial-and-error tests for optimization.(𝐺𝐺, 𝑉𝑉)10403.4 Preliminary results for multi-pass multi-layer simulationPreliminary multi-pass multi-layer simulations are carried out to explore possible grain = structures under different 200 W, ) combinations used in this section are listed in Table 3. The simulation consists of eight layers with a thickness of 200 µm for each layer, and there are six tracks in every layer. The scanning strategy is illustrated in Figure 11; the laser scans in the X direction in a zigzag pattern within each layer (the red arrows) and the scanning direction changes by between two layers. In this work, the XY planes (such as the green plane in Figure 11) are referred as longitudinal sections and the YZ planes (such as the yellow plane in Figure 11) are referred as cross sections.and.

The process parameters used in this section are:= 26 mm/s, and a hatch spacing of 250 µm.𝑁𝑁0𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐𝑉𝑉𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙(𝑁𝑁0, 𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐180°Table 3: Nucleation parameters for section 3.4.Case(m-3)(K)12341013 1014 1014 1015𝑵𝑵𝟎𝟎10 𝜟𝜟𝑻𝑻𝑵𝑵,𝒎𝒎𝒎𝒎𝒎𝒎𝒎𝒎 050Figure 11: The scanning strategy used in section 3.4.First, the grain structures on the center plane of the third track are examined (the grain structures are repeatable for the center planes of different tracks), as shown in Figure 12. It is seen decreases from Figure 12a-d, equiaxed grains become that as increasingly dominant, which is also confirmed in Table 4 with the average aspect ratio of grains decreasing from Figure 12a-d. In Figure 12a-c where large columnar grains are visible, it can be observed that the columnar grains will grow toward the laser scanning direction.

In Figure 12a-c where large columnar grains are visible, it can be observed that the columnar grains will grow toward the laser scanning direction. Since the laser switches the scanning direction between two layers, the growing direction of columnar grains also favors columnar grains the switches between two layers. In Figure 12a where the most ( ) the grain structure is completely columnar; the grains also 𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐 become wider along the build direction. Since there are few new grains introduced, the large columnar grains will only compete with each other; those whose easy-growth directions are better will outgrow the less favored grains. This aligned with the local temperature gradient directionincreases or𝑁𝑁0𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐and𝑁𝑁013−3𝑁𝑁0 = 10, 𝛥𝛥𝑇𝑇𝑁𝑁 = 10𝐾𝐾𝑚𝑚𝐺𝐺⃗1041𝑃𝑃is a typical competitive growth mechanism [12].

It can be foreseen that as more layers are built only a few grains will win the competition and survive, leading to the large columnar grains in Figure 12a. This grain texture is also observed in the experiments [2, 6, 8]. In Figure 10b (), the columnar grains still dominate 𝑁𝑁0 = the grain structure but do not become increasingly wider (and correspondingly the grain size is 10 found to decrease in Table 4). Instead, they are constantly blocked either by new columnar grains (Figure 12b, also experimentally observed in [3]) or a layer of equiaxed grain (Figure 12c, also experimentally observed in [2]). These grain structures can be explained by examining the single- pass grain structures with the same nucleation parameters (in Figure 9a and Figure 8c). In Figure 9a, columnar grains dominate the grain structure with equiaxed grains interspersed among them; when the second layer is built, the top part of the first layer will be re-melted.

In Figure 9a, columnar grains dominate the grain structure with equiaxed grains interspersed among them; when the second layer is built, the top part of the first layer will be re-melted. The interspersed equiaxed grains in the first layer that are partially re-melted will then grow epitaxially to become columnar grains. These “new” columnar grains will join the competitive growth with the existing ones. Since the existing survived grains always have to share the space with the new “competitors”, their growth is limited and the grain size is therefore reduced, as seen in Figure 12b. In Figure 8c, a CET occurs as a thick layer of equiaxed grain at the top of the first layer blocks the epitaxial growth of the columnar grain from the substrate.

In Figure 8c, a CET occurs as a thick layer of equiaxed grain at the top of the first layer blocks the epitaxial growth of the columnar grain from the substrate. When the second layer is built, it cannot re-melt the entire layer of equiaxed grains; the equiaxed grains not re-melted are kept as equiaxed, while the partially re-melted ones act as a substrate from which columnar grains grows epitaxially, leading to the “sandwich” structure observed in Figure 12c. When favors equiaxed grains the most ( ), the equiaxed grain structure is observed in Figure 12d (resembling the experimental observation in [6]), which also has the smallest average 15 𝑚𝑚 grain size and aspect ratio in Table 4.) and 10c (14−315−3, 𝛥𝛥𝑇𝑇𝑁𝑁 = 0𝐾𝐾𝑁𝑁0 = 10, 𝛥𝛥𝑇𝑇𝑁𝑁 = 5𝐾𝐾𝑚𝑚𝑚𝑚and𝑁𝑁0𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐−3𝑁𝑁0 = 10, 𝛥𝛥𝑇𝑇𝑁𝑁 = 0𝐾𝐾Figure 12: Center plane grain structures obtained with different values ofand.

(a):= 10 K, (b):= 1014 m-3,= 0 K, (c):black dashed line indicates the location of fusion line. 𝜟𝜟𝑻𝑻𝑵𝑵𝑵𝑵𝟎𝟎𝜟𝜟𝑻𝑻𝑵𝑵𝑵𝑵𝟎𝟎= 1015 m-3,𝜟𝜟𝑻𝑻𝑵𝑵= 5 K, and (d):𝑵𝑵𝟎𝟎= 1015 m-3,𝜟𝜟𝑻𝑻𝑵𝑵,𝒎𝒎𝒎𝒎𝒎𝒎𝒎𝒎 𝑵𝑵𝟎𝟎𝑵𝑵𝟎𝟎 𝜟𝜟𝑻𝑻𝑵𝑵= 1013 m-3, = 0 K. The1042Table 4 Average grain size and aspect ratio in Figure 12 and Figure 13.average grain size m2)(average aspect ratioFigure 12ab24930 𝜇𝜇 28842.8312.621c3622.239d1991.859a5141.724Figure 13b3622.239c16382.17d9691.92Up to now the examinations of grain structures have been focused on the center planes of the deposition tracks. However, the 3D grain structures in the deposition tracks are more complex than 2D texture observed on the center planes. Case 3 in Table 3 is chosen to demonstrate the different grain structures on different longitudinal sections.

Case 3 in Table 3 is chosen to demonstrate the different grain structures on different longitudinal sections. Figure 13a shows a cross section view of grain structure where three longitudinal sections are marked by the black dashed lines denoted by L1, L2, and L3: L1 is the center plane of a track while L2 and L3 are located away from the center plane towards the edge of the track. Figure 13b-d shows the grain structures on these three longitudinal sections.Figure 13: 3D nature of grain structures. (a): the grain structure on the cross section; L1 is the center plane, L2 and L3 are typical longitudinal sections that are away from the center plane. (b) – (d): the grain structures on the L1, L2, and L3.

(b) – (d): the grain structures on the L1, L2, and L3. The horizontal black dashed line indicates the location of fusion line.It is first found on the cross-section (Figure 13a) that smaller and larger grains alternatively show up along the transverse direction (TD); the grains in the track center (where L1 is located)1043are smaller than those closer to the track edge (where L2 and L3 are located). And on the longitudinal sections (see Figure 13b-d) the grains size are larger on L2 and L3 than in L1 (also see the average grain size in Table 4). The increase of grain size at L2 and L3 can be attributed to a decrease in the occurrence of equiaxed grains. L2 and L3 are located closer to the edge of the tracks, where is usually lower (see Figure 6b and 6c for typical distributions of ) and the occurrence of equiaxed grains is discouraged (see Figure 6a).

𝐺𝐺 Therefore, less new grains are introduced at L2 and L3, and the existing grains become larger 𝐺𝐺 through the competitive growth mechanism. It is also found that the shapes of the grains are different on different longitudinal sections (see Figure 13b-d and the average aspect ratios in Table 4). This can be also attributed to the different grain growth directions in the three longitudinal , the planes. While all the columnar grains tend to grow along the local temperature gradient is different on the three longitudinal sections: when the laser scans along SD in Figure 13a, local the local is aligned very well with SD in L1 and gets increasingly inclined toward the TD in L2 and L3. Therefore, the grains in the three sections stretch along different directions, and the longitudinal sections will cut those grains with different angles with respect to their stretching directions. This causes the grains shapes to be different on the three longitudinal sections.

This causes the grains shapes to be different on the three longitudinal sections. From this demonstration, one can find that analyzing the grain structure simply from one section plane (2-dimensional) can be inadequate to reveal the 3D nature of grain structures.is usually larger and and𝑉𝑉𝑉𝑉𝐺𝐺⃗𝐺𝐺⃗𝐺𝐺⃗and significantly influence the grain structure. Grain structures that are observed in previous experimental works (from columnar-dominated to equiaxed-dominated) are replicated by varying 𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐 nucleation parameters. It is also demonstrated by simulations that the equiaxed grains can block the growth of columnar grains of the previous layer and become new columnar grains in the next layer (Figure 12b and c). Some discrepancies do exist between the simulated grain structures and experimentally observed ones. The equiaxed grains tend to occur near the molten pool tail in simulations, but Ref. [7] reported that equiaxed grains occurred away from the molten pool tail.

The equiaxed grains tend to occur near the molten pool tail in simulations, but Ref. [7] reported that equiaxed grains occurred away from the molten pool tail. Also, fine columnar grains tend to mingle with smaller equiaxed grains (Figure 12b and c) in the simulations, due to the constant nucleation ahead of the SF; but some experimental results can have only the fine columnar grains without too many equiaxed grains visible [7, 35]. These discrepancies can be attributed to multiple reasons such as different thermal history, materials, nucleation conditions and/or mechanisms. Therefore, experimental calibrations and validations about the current model need to be conducted and are considered as important future works.It can be concluded from this preliminary study that nucleation parameters𝑁𝑁04.

SummaryIn this work, the thermal history of a typical MAM process, laser direct deposition, is simulated by a finite volume model, and then cellular automata method combined with a heterogeneous nucleation model is used to simulate the grain structure. Single-pass simulations , and are first conducted to demonstrate the effects of nucleation parameters, process parameters, , on the grain structure. After that, preliminary explorations of the grain structure of a multi-pass multi-layer process are made using different nucleation parameters.

After that, preliminary explorations of the grain structure of a multi-pass multi-layer process are made using different nucleation parameters. 𝑉𝑉𝑙𝑙𝑙𝑙𝑙𝑙𝜌𝜌𝑙𝑙 The major findings of this work are summarized as follows:andand𝑁𝑁0𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐𝑃𝑃1044Larger nucleation densitywill encourage the heterogeneous nucleation ahead of the solidification front and lead to equiaxed- favored grain structures.and smaller critical nucleation undercooling𝑁𝑁0 and scanning velocity and solidification front velocity𝛥𝛥𝑇𝑇𝑁𝑁,𝑚𝑚𝜌𝜌𝑙𝑙𝑐𝑐Laser power will significantly influence the temperature gradient on the fusion surface, which can determine the grain structure being completely columnar, completely equiaxed or mixed with both columnar 𝐺𝐺 and equiaxed grains. 𝑃𝑃Laser power will significantly influence the temperature gradient on the fusion surface, which can determine the grain structure being completely columnar, completely equiaxed or mixed with both columnar 𝐺𝐺 and equiaxed grains.

𝑃𝑃columnar grains, leading to complex varying grain structures.The grain structures are 3D in nature; they can vary significantly when examined on differentplanes. Important future works include to validate the modeled nucleation and growth mechanisms, as well as to calibrate the nucleation parameters through experiments.Reference[1] W. E. Frazier, "Metal additive manufacturing: a review," Journal of Materials[2][3][4]Engineering and Performance, vol. 23, no. 6, pp. 1917-1928, 2014. L. L. Parimi, G. Ravi, D. Clark, and M. M. Attallah, "Microstructural and texture development in direct laser fabricated IN718," Materials Characterization, vol. 89, pp. 102-111, 2014. G. Dinda, A. Dasgupta, and J. Mazumder, "Evolution of microstructure in laser deposited Al–11.28% Si alloy," Surface and Coatings Technology, vol. 206, no. 8, pp. 2152-2160, 2012. V. Ocelík, I. Furár, and J. T. M.

Desbiolles, and P. Thevoz, "Prediction of grain structures in various solidification processes," Metallurgical and Materials Transactions A, vol. 27, no. 3, pp. 695-705, 1996.[34] T. T. Roehling et al., "Modulating laser intensity profile ellipticity for microstructural control during metal additive manufacturing," Acta Materialia, vol. 128, pp. 197-206, 2017.[35] G. Dinda, A. Dasgupta, S. Bhattacharya, H. Natu, B. Dutta, and J. Mazumder,"Microstructural characterization of laser-deposited Al 4047 alloy," Metallurgical and Materials Transactions A, vol. 44, no. 5, pp. 2233-2242, 2013.1047.

DIMENSIONAL FINITE ELEMENT MODELING OF SELECTIVE LASER MELTING TI-6AL-4V ALLOYC.H. Fu, Y.B. Guo Dept. of Mechanical Engineering, The University of Alabama, Tuscaloosa, AL 35487, USAAbstractSelective laser melting (SLM) is widely used in making three-dimensional functional parts layer by layer. Temperature magnitude and history during SLM directly determine the molten pool dimensions and surface integrity. However, due to the transient nature and small size of the molten pool, the temperature gradient and the molten pool size are very challenging to measure and control. A 3-dimensional finite element simulation model has been developed to simulate multi-layer deposition of Ti-6Al-4V in SLM. A physics-based layer build-up approach coupled with a surface moving heat flux was incorporated into the modeling process. The melting pool shape and dimensions were predicted and experimentally validated.

The melting pool shape and dimensions were predicted and experimentally validated. Temperature gradient and thermal history in the multi-layer build-up process was also obtained. Furthermore, the influences of process parameters and materials on the melting process were evaluated.Keywords: Selective laser melting, FEA, Temperature gradient, Molten pool1. Introduction1.1. Selective laser melting (SLM)Layer manufacturing (LM) has been developed for over 20 years. It is widely used in making three-dimensional parts directly from CAD models by adding material layer by layer, rather than removing material. Rapid manufacturing (RM), as a layer manufacturing process, is aiming at producing functional parts that can be directly used as end-use products rather than prototypes.

Rapid manufacturing (RM), as a layer manufacturing process, is aiming at producing functional parts that can be directly used as end-use products rather than prototypes. The advantages of RM include: (1) it allows fabrication of geometrically complex parts, even for parts that not possible to fabricate with subtractive manufacturing processes [1] and (2) it is a cost-effective process for single parts and small batches. The current challenge for RM is that it should guarantee long-term consistent components that have required physical, mechanical, and geometrical properties [2].Selective laser melting (SLM) is a typical powder-based RM processes. SLM features fully melting of powder, thus near full density parts can be produced. The advantage of SLM over electron beam melting (EBM) and laser metal deposition (LMD) is that the equipment is relatively less complex. There is no need for high vacuum chamber to guide electron beam as opposed to EBM.

There is no need for high vacuum chamber to guide electron beam as opposed to EBM. There is also no stringent alignment needed between powder feeding nozzle and laser beam (as opposed to LMD) [3]. SLM is widely used in aerospace, automotive, biomedical, and energy industries [4-8]. Corresponding author. Tel.: +1 205 348 2615; Fax: +1 205 348 6419. E-mail address: yguo@eng.ua.edu (Y.B. Guo).Even though SLM is a very promising technique, due to the intense heat input, problems such as balling effect, deteriorated surface finish, tensile residual stress, and part distortion are still frequently observed in SLM process.1.2. Process characteristicsFig. 1(a) shows the schematic of SLM process. It uses a fine powder system to distribute a 20-100 µm layer onto the substrate plate inside a chamber containing inert gas. Then the layer is fused by selectively scan laser beam on the powder surface. The intensive laser energy fully melts the metal powders to form a solid metal.

Then the layer is fused by selectively scan laser beam on the powder surface. The intensive laser energy fully melts the metal powders to form a solid metal. This process is repeated layer after layer until the part is complete.Laser systemYXCooling AX YHeating ALaser beamLaser windowChamberThermal contraction YAσyLaser beamThermal expansion AY σyPowder systemA-AA-ASolidified zoneMolten poolPowder layerSLM partHAZRetractable platformSubstrate(a)(b)Fig. 1 Selective laser melting experimental setup and process principle.The process principle of SLM is shown in Fig.

1 Selective laser melting experimental setup and process principle.The process principle of SLM is shown in Fig. 1(b), when laser beam scans on the powder surface, energy is transferred from the top surface to subsurface through various physical phenomena, such as absorption and scattering of the laser radiation, heat transfer, fluid flow within the molten pool, melting/evaporation, and chemical reactions.Powder melting happens when the temperature reaches the material melting temperature. The dimension of the molten pool is greatly affected by the average applied energy per unit volume, which is mainly controlled by process parameters including laser power and scanning speed. It has been reported that higher scanning speed leads to an increasing length to diameter ratio in a molten pool. Such a molten pool with length to diameter ratio greater than π will cause balling effect [9].

Such a molten pool with length to diameter ratio greater than π will cause balling effect [9]. The balling effect can be characterized by the break-up of the molten pool into balls, which degrades surface integrity of the final components. Therefore, to achieve the optimal surface integrity, the range of the optimal scanning speed is limited by the balling effect. It has also been reported that a higher laser power can expand the range for the optimal speed [10].The nature of laser heating of a powder layer is considerably different from laser heating of a solid bulk body because of two reasons. First, laser absorptivity for powder layers is substantially higher than that of the bulk body due to their different granulomorphometry and density [10,11].Second, thermal conductivity of powder bed is much lower than that of the opaque body since powder particles have limited contact area with each other, thus, heat transfer is only dominated by the low thermal conductivity of gas [12].

Due to these dissimilarities, it can be inferred that changing powder diameters and compactness can significantly affect heat conduction property during SLM.In a SLM process, a very steep temperature gradient is generated due to the rapid heating of the upper surface by laser and slow heat conduction of underlying layers [9]. Temperature gradient induced surface tension gradient in the molten pool can cause convective motions, known as the Marangoni flow. Temperature gradients also induce high internal stresses. As shown in Fig. 1(b), when laser beam hits the top surface, the expansion of the heated layer is restricted by the vicinity area, therefore, a compressive stress is generated on the top surface. When the yield stress of the material is reached, plastic distortion occurs. As the laser moves away, cooling takes place; the contraction of the top layer is then restricted by the surrounding area, leading to a tensile residual stress on the top surface.

As the laser moves away, cooling takes place; the contraction of the top layer is then restricted by the surrounding area, leading to a tensile residual stress on the top surface. The tensile residual stress will accumulate during each laser scan, which can cause delamination or cracks. Also, the cooling and shrinkage will create shorter top layers than the bottom layers, thus, the component is distorted by bending towards the laser beam.Products are formed from the consolidated material in SLM process. The temperature gradient controls the formation, motion, and rapid solidification of the molten pool, which directly determines the microstructure, material properties, and final surface integrity of the components. It was found that the material processed by SLM exhibit a very fine, non- equilibrium microstructure [13].

It was found that the material processed by SLM exhibit a very fine, non- equilibrium microstructure [13]. However, the excessive heat input will elongate the grains and induce an heat affected zone (HAZ), which will degrade the material property and performance.The temperature gradient mechanism is critical for SLM due to the following reasons: it directly controls the dimension and stability of a molten pool, which plays a large role in determining surface roughness and porosity of the parts. Temperature gradient also controls the formation of the HAZ, tensile stresses, and part distortion. If temperature gradient can be accurately measured and controlled, surface integrity and functionality of the final parts can be significantly improved. However, the temperature gradient is very challenging to measure due to its very small size and transient nature. Therefore, a finite element simulation was used in this study to gain insight into the process mechanism.

Therefore, a finite element simulation was used in this study to gain insight into the process mechanism. The goal of this study is to understand the fundamental temperature gradient mechanism by developing a finite element model.1.3. Literature review on SLM simulationA few research has been conducted to simulate an SLM process to predict temperature gradient, stress formation, and part distortion. A successful finite element model of SLM needs to consider four coherent aspects: (1) a thermal model representing scanning laser thermal input, (2) a material model to define the temperature-dependent material property in the cyclic heat and cooling environment, (3) a physics-based method to describe the layer build-up process, and (4) a laser material interaction model that accounts for the transient thermomechanical phenomena.A thermal model with spatial accuracy is critical to simulate the laser heating in SLM.

It has been proved that thermophysical properties such as thermal conductivity and specific heat change significantly when material is heated up to liquid phase from solid phase [19]. Moreover, the thermal conductivity for solid substrate and powder bed are also significantly different. The effective thermal conductivity of loose metallic powders is controlled by gas in the pores, and it was found that for powders of 10-50 µm diameter, the effective thermal conductivity is typically from 0.1-0.2 W/mK. However, the thermal conductivity starts to rise sharply when the material starts to melt. Therefore, it is critical to include these property changes in SLM simulation.To simulate the layer build-up process in SLM, Ding et al. [20] used multiple sequential steps in the simulation. More specifically, after each scan, the temperature from the previous step was imported into the next step for newly added layers. Another technique is to use the element “birth and death”.

More specifically, after each scan, the temperature from the previous step was imported into the next step for newly added layers. Another technique is to use the element “birth and death”. The deposition of powder material on the powder bed to create a solid is analogous to the activation of new elements at desired time point [14]. Dai et al. [21] simulated the layer build-up process in SLM by adding all the elements within one layer simultaneously onto the previous layer. After adding the entire layer of the powder elements, the heat flux was applied on this newly added layer to simulate the laser thermal processing.Although these simulation studies have explored the simulation of SLM process on different materials using different thermal modeling and element “death/birth” methods, little research has been done on a systematic finite element simulation of SLM including a moving surface heat flux, realistic layer build-up, and temperature-dependent multi-phases material properties.

SLM is a very complex process with wide arrange of transient non-equilibrium phenomena taking place. Not only laser parameters, such as laser power or laser spot size, but also the materialRef.[14][15,16][17][18]properties such as temperature-dependent thermal properties play an important role in determining temperature gradient mechanism and molten pool behavior. Therefore, the objectives of this study are to: (1) develop a 3-dimensional finite element simulation model of SLM with an emphasis on a physics-based layer build-up approach coupled with a surface moving heat flux, (2) predict molten pool shape and dimensions with experimental validation, (3) study the effect of process parameters and materials on the melting process, and (4) gain insights into the temperature gradient mechanism of SLM process.2. Simulation Approach2.1.

Simulation Approach2.1. Model descriptionThe finite element analysis (FEA) package Abaqus/Standard was used to incorporate both the moving heat flux subroutine and material properties to simulate the laser-material interaction in SLM. The mesh design is shown in Fig. 2. The workpiece is divided into two sections: substrate and powder layers. The dimensions of the substrate are 4 mm (length) × 1 mm (width) × 0.5 mm (thickness). The dimensions of the powder layer are 2 mm (length) × 0.2 mm (width) × 0.15 mm (thickness). The layer thickness is 30 µm and five layers are built-up in the simulation. The laser scanning direction is along the X-axis. Element size is biased with a higher density in the powder layers and top of the substrate. Within the fine mesh, the element size is 10 µm (length) × 5 µm (width) × 5 µm (thickness). The boundary condition on the bottom plane is fixed to provide proper constraint of the workpiece.

Within the fine mesh, the element size is 10 µm (length) × 5 µm (width) × 5 µm (thickness). The boundary condition on the bottom plane is fixed to provide proper constraint of the workpiece. Also, the model is symmetric with respect to X-Z plane so that only half of the workpiece needs to be modeled to decrease the computational time. The initial temperature is set to room temperature (20 °C).10200150500Layers555 layersLayers2000XLayer thickness: 30 µmSubstrate4000Z Unit: µmY1000Fig. 2 Simulation schematic of SLM.The simulation was performed using Abaqus/Standard since the moving heat flux subroutine DFLUX can be programed with the implicit solver. Another advantage of using the implicit solver is that the temporal discretization is more stable despite a certain reduction in computational efficiency.2.2.

Second, all the elements within the powder layers are deactivated at the beginning of the analysis. Last, the elements in the first powder layer are activated followed by the first laser scan. This process is repeated for the successive layers so that a very close approximation of the actual SLM process can be simulated, where powders are laid layer by layer after each scan. Since the powder layer thickness is 30 µm in simulation and each powder layer contains 5 layers of elements for high mesh resolution, there are 48,000 elements for each powder layer. The powder depth in the experiment is the same as the powder layer thickness in simulation. A schematic of the layer build-up process is shown in Fig. 4.(2)Powder meshScan of 1st layerScan of 2nd layerDeactivated elementsActivated elementsFig. 4 Schematic of layer build-up process.2.4. Material propertiesTi-6Al-4V was used in this study since it is widely used in SLM for aerospace and biomedical applications.

4 Schematic of layer build-up process.2.4. Material propertiesTi-6Al-4V was used in this study since it is widely used in SLM for aerospace and biomedical applications. The temperature-dependent material thermal properties of Ti-6Al-4V powder and substrate are listed in Table 2.

Simulation conditionsThe simulations were a sensitivity analysis to determine the effects of laser power on molten pool geometry and temperature gradient. The effect of material property, powder vs. solid, was also studied. Two different material properties were used for the elements in the build-up layers. In the first set of simulations, the powder properties with a low thermal conductivity were used, while the solid properties with a high thermal conductivity were used in the second set of simulations. This approach would provide insights of the effect of powder compactness on the pool geometry and temperature gradient mechanism.The laser in this study is a continuous Nd:YAG laser (wavelength = 1.06 µm) that is widely used in actual SLM process. At this wavelength, the absorption coefficient for Ti-6Al-4V powder was assumed to be the same as pure titanium powder, which is equal to 0.77 [11]. The laser scanning speed is 0.2 m/s.

At this wavelength, the absorption coefficient for Ti-6Al-4V powder was assumed to be the same as pure titanium powder, which is equal to 0.77 [11]. The laser scanning speed is 0.2 m/s. The laser spot radius is 26 µm and the layer thickness is 30 µm for all the simulation cases.3. Model Validation and DiscussionsFig. 5(a) shows the representative temperature contour when laser was scanning on the top layer. The molten pool dimension is indicated by the gray region, where the elements within this space have a temperature above the melting temperature of Ti-6Al-4V. A “comet tail” profile can be seen in Fig. 5(b). This asymmetry can be attributed to the movement of the laser as well as the fact that the heated-up material has greater conductivity than the untreated powder in front of the laser.

5(b). This asymmetry can be attributed to the movement of the laser as well as the fact that the heated-up material has greater conductivity than the untreated powder in front of the laser. This skewed temperature distribution towards the rear of the laser was also reported in other studies [14,16].TemperatureL(melting length)Laser center(melting width)WMolten poolYZX(cid:1848) (cid:3404)(cid:2024)(cid:1830)(cid:1838)(cid:1849) 6(melting depth)DMolten pool geometry(a)(b)Fig. 5 (a) Representative temperature contour and (b) molten pool geometry.3.1. Molten pool geometry and dimensionsMelting depth: The remelting depth for each layer has proven to be critical to microstructure evolution [13], remove contaminants [24], and prevent balling effect [10]. Therefore, it is important to monitor, predict, and control the melting depth for each layer during SLM.Fig. 6 shows the effect of laser power on melting depth.

Therefore, it is important to monitor, predict, and control the melting depth for each layer during SLM.Fig. 6 shows the effect of laser power on melting depth. It can be seen that as the laser power increased from 20 W to 80 W, the predicted melting depth increased from 20 µm to 40 µm, which is larger than the powder layer thickness. Remelting started to happen at laser power of 40 W. Moreover, as illustrated by the error bar, the predicted melting depth for all the five layers at the same laser power were very close, which indicates that the preheating of the previous layers had little influence on remelting of the subsequent layers at the concerned conditions.120EXP)m µ (100SIM_PowderSIM_Solid)D80( h p e d g nti t l6040e M200020406080100Laser power (W)Fig.

6 Effect of laser power on melting depth.To validate the simulation results, the predicted melting depths were compared with the experimental data at same SLM conditions [22]. The SLM experiments were conducted with a powder layer thickness of 30 μm by building a single track on a solid substrate. For comparison purpose, the simulation conditions are same as the experimental conditions. The predicted width and depth of molten pool were analyzed by optical microscopy based on the solidified microstructure. It can be seen that the predicted data agreed with the experimental data with in trend. The measured melting depths in experiments were higher than the predicted ones.

It can be seen that the predicted data agreed with the experimental data with in trend. The measured melting depths in experiments were higher than the predicted ones. The assumptions and error sources responsible for the discrepancies are as follows: (1) the absorptivity, specific heat, and thermal conductivity of the powder bed were obtained from the literature, which may be different from the actual values since these properties are greatly affected by powder size, compactness, and surrounding environment; (2) there may be uncertainties in experimental inputs that differentiate the experimental conditions with simulation, e.g., the laser power loss during delivery, laser spot size change due to defocus etc.; (3) there are measurement errors as indicated by the large size of the error bars.

In addition, the experimental measurements were based on the microstructural boundaries are often difficult to characterize and could induce high non- uniformity and randomness in the measurement. These compounding factors influence the predictive accuracy of the simulation approach, which can be improved in the future.the solidified microstructure, whereMelting length/width: The predicted melting width is shown in Fig. 7. The predicted melting depth agreed with the experimental data in trend. As the laser power increased, the melting width almost increases linearly. The predicted melting length is shown in Fig. 8. For all the cases, the predicted length to width ratio is near 1.3, which is smaller than the critical value of π for balling effect [9]. This indicates that there was no balling effect at concerned conditions.200EXPSIM_Powder)m µ (150SIM_Solid)W( h t d w g n i t lie M100500020406080100Laser power (W)Fig.

7 Effect of laser power on melting width.80SIM_Powder)m µ (60SIM_Solid) L ( h g n et40lg ni t le M20002060 Laser power (W)4080100Fig. 8 Effect of laser power on melting length.Molten pool volume: It can be seen in Fig. 5 that the molten pool was in a near ellipsoidshape, and the volume of the molten pool can be approximated by(cid:1848) (cid:3404)(cid:2024)(cid:1830)(cid:1838)(cid:1849) 6(3)where D	is the melting depth, L is the melting length, W is the melting width of the molten pool, respectively. As indicated in Fig. 9, when laser power increased, the molten pool volume increased dramatically.8E+05SIM_Powder) 3m µ ( e m u o vl6E+054E+05SIM_Solidlo o p n e t l o M2E+050E+00020406080100Laser power (W)Fig. 9 Effect of laser power on molten pool volume.The effect of material property: The changes of the molten pool depth, width, and volume were also plotted in Figs. 6-9.

9 Effect of laser power on molten pool volume.The effect of material property: The changes of the molten pool depth, width, and volume were also plotted in Figs. 6-9. It can be observed that by using the solid property (higher thermal conductivity) in powder layers, the molten pool depth, width, length, and volume decreased. This finding suggests that the higher thermal conductivity of the solid “delivered” the thermal energy into the deeper and further vicinity, thus steep temperature gradient caused by the very low thermal conductivity of the powder was reduced. Therefore, it can be inferred that by using a more compact powders, steep temperature gradient can be reduced and the molten pool size can be decreased.3.2. Temperature gradientDepth direction: Temperature distributions in the subsurface are shown in Fig. 10. The temperatures decrease dramatically from the laser beam’s center to the subsurface.

Temperature gradientDepth direction: Temperature distributions in the subsurface are shown in Fig. 10. The temperatures decrease dramatically from the laser beam’s center to the subsurface. At a laser power of 20 W, the temperature difference within 100 µm was smaller when compared to other 3 simulation cases (40 W, 60 W, and 80 W). As the laser power increases, temperature gradient becomes steeper. In addition, within 40 µm depth, the temperature distribution was significantly different for the concerned simulation cases. Beyond 40 µm, the temperature distribution for the simulated cases shared similar magnitude. This was due of the fact that at higher laser powers, more thermal energy was accumulated near the top surface due to the low thermal conductivity. Also, extreme high temperature was observed at the simulated conditions, this is non-practical since material would evaporate at such high temperatures.1600012000Laser power 20 W 40 W 60 W 80 WNode path)C ° (.

12. The temperatures decrease as moving away from the laser beam’s center along the scanning direction. When compare the temperature gradients in depth, width, and laser scanning direction, it can be observed that the temperature gradient was much steeper in the depth direction. It is due to that the thermal conduction of powders in depth direction wasconstrained by the thermal conductivity rather than the direct laser radiation in the width and laser scanning directions.16000Laser power12000Node path20 W 40 W 60 W 80 W)C ° (. p m e T8000Scan speed: 0.2 m/s Laser radius: 26 µm Layer thickness: 30 µm40000020406080100Distance to laser center (µm)Fig. 12 Temperature gradient in laser scanning direction.Thermal history: The thermal history at the selected location is shown in Fig. 13. The selected point was located on the top surface and in the middle of the first layer.

13. The selected point was located on the top surface and in the middle of the first layer. The peak temperatures represent the start of each heating-cooling cycle during each laser scan. As seen in the figure, the material point experiences five heating-cooling cycles. At the lowest laser power (20 W), the 1st layer was raised above its melting temperature only once during the first laser scan, while at the highest laser power (80 W), the 1st layer was melted 3 times. As more layers were deposited in the 4th and 5th laser scan, the temperature drops below the melting temperature. This suggests that addition of layers and subsequent scanning are of great importance to the previous layers.

This suggests that addition of layers and subsequent scanning are of great importance to the previous layers. It can be further noticed that cooling to the ambient temperature after each laser scan occurred only within a few milliseconds, which indicating the material was subjected to rapid heating-cooling cycles, which leads to very high tensile residual stresses [25].16000scan 1Center point on layer top in 1st scanLaser power 20 W 60 W40 W 80 W1500scan 312000scan 4)C ° (1000scan 5. p m e T80005004000scan 20 0.020.030.040.05scan 3scan 4scan 5000.010.020.030.040.05Time (s)Fig. 13 Temperature history of center point on layer top in 1st scan.3.3. RecommendationThe developed finite element model showed certain predictive capabilities.

13 Temperature history of center point on layer top in 1st scan.3.3. RecommendationThe developed finite element model showed certain predictive capabilities. To improve the predictive accuracy, it is recommended that the absorptivity, specific heat, and thermal conductivity of the powder bed should be measured experimentally and used in simulation. Moreover, some other factors such as power loss, spot size, etc. should be considered in future research.Another important factor that should be considered in the future is the computational time. More specifically, for each micro-scale simulation case in this study, the computational time was 240 hours with a standard desktop workstation. It is estimated that for a practical macro-scale selective laser melting part, the computational time would be a million year if the same mesh size is used, which is not practical.

It is estimated that for a practical macro-scale selective laser melting part, the computational time would be a million year if the same mesh size is used, which is not practical. Therefore, it is critical to develop a scaling-up methodology from micro-scale level to macro-scale simulations without a significant cost of computational time. In this way, real parts with residual stress and distortion could be predicted.4. ConclusionsThis study focuses on the development of a novel simulation model for SLM. The effects of laser powder and material properties on molten pool geometry and temperature gradient were assessed. The predictions of molten pool geometry were verified by the experimental data. Key findings are summarized as follows:A new model of a physics-based layer build-up approach coupled with a surface moving heat flux was developed to simulate SLM.The molten pool geometries was predicted and verified by the experimental data with reasonable accuracy.

It was predicted that balling will not occurs under the tested condition. It was also predicted that by using compact powders, the steep temperature gradient can be reduced and the molten pool size can be decreased.The temperature gradient in the depth direction is much higher than those in width and laser scanning directions since the heat transfer in the depth direction is greatly constrained by the low thermal conductivity of the powder layers. The addition of subsequent layers significantly affected the temperature gradient and the melting of previous layers.

The addition of subsequent layers significantly affected the temperature gradient and the melting of previous layers. It only takes few milliseconds to melt the material and cool it to room temperature, which indicated the very fast heating and cooling cycles in SLM.AcknowledgementThe first author would like to thank the Graduate Council Fellowship at the University ofAlabama for the financial support.References [1] Waterman, N.A., Dickens, P., 1994, "Rapid Product Development in the USA, Europeand Japan," World Class Design to Manufacture, 1(3), pp. 27-36.[2] Levy, G.N., Schindel, R., Kruth, J.P., 2003, "Rapid Manufacturing and Rapid Tooling with Layer Manufacturing (LM) Technologies, State of the Art and Future Perspectives," CIRP Annals - Manufacturing Technology, 52(2), pp. 589-609.[3] Guo, N., Leu, M., 2013, "Additive Manufacturing: Technology, Applications andResearch Needs," Frontiers of Mechanical Engineering, 8(3), pp.

589-609.[3] Guo, N., Leu, M., 2013, "Additive Manufacturing: Technology, Applications andResearch Needs," Frontiers of Mechanical Engineering, 8(3), pp. 215-243.[4] Vandenbroucke, B., Kruth, J., 2007, "Selective Laser Melting of Biocompatible Metals for Rapid Manufacturing of Medical Parts," Rapid Prototyping Journal, 13(4), pp. 196- 203.[5] Clare, A., Chalker, P., Davies, S., Sutcliffe, C., Tsopanos, S., 2008, "Selective Laser Melting of High Aspect Ratio 3D Nickel-Titanium Structures Two Way Trained for MEMS Applications," International Journal of Mechanics and Materials in Design, 4(2), pp. 181-187.[6] Hollander, D.A., von Walter, M., Wirtz, T., Sellei, R., Schmidt-Rohlfing, B., Paar, O., Erli, H., 2006, "Structural, Mechanical and in Vitro Characterization of Individually Structured Ti–6Al–4V Produced by Direct Laser Forming," Biomaterials, 27(7), pp.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed Paper316L POWDER REUSE FOR METAL ADDITIVE MANUFACTURINGB. Sartin*, T. Pond*, B. Griffith*, W. Everhart*, L. Elder*, E. Wenski*, C. Cook*, D. Wieliczka*, W. King†, A. Rubenchik†, S. Wu†, B. Brown*, C. Johnson*, and J. Crow*Honeywell Federal Manufacturing and Technology under Contract DE-NA0002839 †Lawrence Livermore National Laboratory under Contract DE-AC52-07NA27344AbstractMetal additive manufacturing via laser powder bed fusion is challenged by low powder utilization. The ability to reuse metal powder will improve the process efficiency. 316L powder was reused twelve times during this study, completing thirty-one builds over one year and collecting 380 powder samples. The process, solidified samples, and powder were analyzed to develop an understanding of powder reuse implications.

The process, solidified samples, and powder were analyzed to develop an understanding of powder reuse implications. Solidified sample characteristics were affected more by slight process variations than by cycling of the powder. While a small percentage of powder was greatly affected by processing, the bulk powder only observed a slight increase in powder size.IntroductionThe effects on metal powder from processing through laser and electron beam powder bed additive manufacturing (AM) systems are key concerns to industry. An understanding of the implications of powder reuse on the final part characteristics must be developed. This understanding is necessary to ensure component quality while enabling reduction of fabrication costs through powder reuse.Powder reuse research has been reported by many, especially for Ti64 and nickel alloys [1 - 14]. Typically, these studies found powder particle size to increase [1 - 5] and oxygen levels to rise [4 - 10] with reuse.

Typically, these studies found powder particle size to increase [1 - 5] and oxygen levels to rise [4 - 10] with reuse. When mechanical testing was performed, strength increased and elongation reduced in correspondence with the increase in oxygen levels [8, 5]. Limited work has been performed on steels, including 17-4SS [15, 16], 316L [17], and 4340 [18]. Slotwinski et al. found particle size to increase [15] while others documented no consistent trend in size or morphology [16]. Skryecki is the only 316L powder reuse study reviewed, finding only minor variations in particle size and morphology [17]. The reduction in fines in the powder distribution and an increase in larger powder generated through processing were found to improve flowability as determined by various methods [1, 4, 13, 16, 19]. Generally, every finding has a conflicting conclusion in another study, with numerous authors finding no significant changes in the powder or solidified samples.

Generally, every finding has a conflicting conclusion in another study, with numerous authors finding no significant changes in the powder or solidified samples. Most authors conclude practical powder reuse may be used with minimal impact on the process results.In this powder reuse study, a 500kg single lot of 316L stainless steel was processed for twelve reuse cycles.

The study was executed over one year, thirty-one builds, and yielded pertinent information on metal AM powder, process, and equipment; and continued to advance the methods for analyzing metal AM powder.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing and Technology, LLC under Contract DE-NA0002839 and Lawrence Livermore National Laboratory operated and managed by Lawrence Livermore National Laboratory , LLC under Contract DE-AC52-07NA27344351ExperimentalPowder used by this study was processed through a Renishaw AM250 laser powder bed AM machine and ancillary sieve according to the process displayed in Figure 1. Powder samples (~500g) are removed at each processing point and solid samples are removed for testing after each build.

Powder samples (~500g) are removed at each processing point and solid samples are removed for testing after each build. Powder was processed through each reuse cycle collectively to maintain discrete cycles of powder over the course of multiple builds as presented in Figure 2.Sieve Debris removedSolidified Samples removedVirgin powder pre-sieveVirgin powder post-sieve1 Cycle powder pre-sieveFigure 1: Approach flow chartFigure 2: Representative build layout of solidified samplesBuild file and parameter definition processing information are presented in Table 1. In addition to these easily controlled inputs, a best effort was made to control other aspects of the process including powder storage along with setup and operation of the equipment (AM machine and sieve).

Tests used include but were not limited to combustion or inert gas fusion (LECO), acid digestion followed by iCAP, direct insertion probe mass spectroscopy, ASPEX SEM, Rotap sieving, Helium pycnometry, Karl Fischer titration, vapor sorption, X-ray photoelectron spectroscopy (XPS), scanning electron microscope (SEM) and energy dispersive X-ray spectroscopy (EDS), X-ray diffraction (XRD), calorimetric absorptivity determination [20, 21], powder rheology, Hausner ratio, and Scott volumeter.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing and Technology, LLC under Contract DE-NA0002839 and Lawrence Livermore National Laboratory operated and managed by Lawrence Livermore National Laboratory , LLC under Contract DE-AC52-07NA27344353Samples fabricated during each build included tensile bars and density pucks.

The as-AM tensile test specimens fabricated, corresponding to those used in prior work [22], were similar to that presented in ASTM E8M as subsize R4, with the exception of a 1° gage taper as opposed to a 1% max increase in gage diameter. These samples were produced and tested with an as-AM surface finish. Solidified samples produced from various cycles of reused powder were tested to analyze composition, corrosive properties, crystalline structure, density, porosity, microstructure, mechanical strength and ductility, and surface finish.

Tests used include but were not limited to combustion or inert gas fusion (LECO), acid digestion followed by iCAP, copper sulfate testing, XRD, density via Archimedes in alcohol, metallography, Knoop microhardness, tensile testing, SEM fractography, and surface profilometry.Results and DiscussionProcessingOperating metal AM equipment for an extending period reveals numerous gaps in the current processing equipment and procedures when documented with the rigor associated with this effort. The metal AM equipment used by this study required a significant amount of operator intervention to continue operations as displayed in Figure 6. Despite consistent build geometries, the build duration ranged from 6 to 43 days with an average build duration of 10.4 days (σ = 7.9).

Despite consistent build geometries, the build duration ranged from 6 to 43 days with an average build duration of 10.4 days (σ = 7.9). Continued maturation of metal AM equipment will reduce these challenges and allow for more consistent processing.Figure 6: Build event countsFigure 5: Study mass averages by collection ptPowder mass values were recorded for all builds to document the percentage of powder associated with each equipment collection location (Figure 3 and Figure 4) and the amount consumed by each processing step. This study found that on average 6.7% of the powder introduced to the system was consumed by the metal AM process when confined to the metal AM equipment, including approximately 2-3% consolidated into the components fabricated, ~1% collected in the AM equipment filtration, and the remainder consumed during the AM equipment and component clean-out process.

An additional 3% of the powder was removed during the sieving ©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing and Technology, LLC under Contract DE-NA0002839 and Lawrence Livermore National Laboratory operated and managed by Lawrence Livermore National Laboratory , LLC under Contract DE-AC52-07NA27344354process and ~3.5% of the powder was removed as samples to facilitate powder characterization at each cycle of reuse. This process powder utilization of 2-3% aligns with the 3% utilization reported by Hann [10] and the 3 - 6% utilization reported by Kellens et al. [23]PowderInorganic analysis was performed on 228 samples of powder to characterize variation based on collection location and cycle of reuse. No statistical variation was found between collection locations (powder bed, front overflow, back overflow) which will enable reduced sampling going forward.

No statistical variation was found between collection locations (powder bed, front overflow, back overflow) which will enable reduced sampling going forward. Powder chemistry is affected by the metal AM process; powder collected from the 80µm sieve screen in-between each reuse cycle (debris) tested considerably higher for oxygen and slightly higher for hydrogen (Figure 7). Additionally, the distribution of composition was widened on these samples for oxygen, hydrogen, carbon, sulfur, and titanium. The removal of debris via sieving removed the vast majority of powder that had been chemically altered, however, as there were no statistical trends in the elemental composition observed relative to reuse cycle (Figure 8).Figure 7: Histograms by processing step, selected test dataFigure 8: Interval plots of selected elements vs.

powder reuse cycleOrganic testing was performed via direct probe mass spectroscopy on a number of articles that comprised the sieving equipment, metal AM equipment, or are otherwise in the process and considered opportunities for contamination. Notable findings include dioctyl phthalate that is free to migrate from the polyvinyl chloride tubing used on the AM equipment and sieving equipment; and acetic acid residue potential from the application of vinegar based WindexTM in the equipment cleaning procedure.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing and Technology, LLC under Contract DE-NA0002839 and Lawrence Livermore National Laboratory operated and managed by Lawrence Livermore National Laboratory , LLC under Contract DE-AC52-07NA27344355296 powder samples were characterized for size via ASPEX SEM and Rotap sieving.

No statistically significant change in size or shape was observable via ASPEX SEM. Particle size distribution as observed via Rotap sieving identified a statistical difference between virgin and eight cycle powder (seven cycles of reuse) with a decrease in the percentage volume of 20 to 25µm powder and an increase in the percentage volume of 45µm and larger powder.Figure 9: Reused post-sieve powder (combined all collection groups) vs. virgin by cycleNo statistical change was observed for a number of characteristics tested including skeletal density, flowability, and moisture. Skeletal density was performed via Helium pycnometry and remained consistent over eight cycles; measured at 7.957 +/-0.024 g/cc on virgin powder and 7.964 +/-0.030 g/cc on eight cycle powder. Flowability was determined via various methods including Scott’s volumeter apparent density and Hausner ratio.

Flowability was determined via various methods including Scott’s volumeter apparent density and Hausner ratio. The test variability associated with these tests did not allow for differentiation between flowability of various cycles of powder. Moisture was characterized by Karl Fisher titration and vapor sorption. Both methods observed low moisture adsorption in powder, never exceeding 0.0092% by weight.Flowability testing performed using an MCR302 two vane stirrer at 8 rpm, did differentiate between virgin, four cycle, and eight cycle powder. Three 80 mL samples were tested following air fluidization in the test cell at 1.6L/min. Values reported as cohesion strength were reported as 608, 681, and 771 Pa for the virgin, four cycle, and eight cycle samples, respectively.Surface characterization was performed on powder after various cycles of reuse via several methods.

XPS was used on numerous powder samples with example surveys displayed in Figure 10, where red is as-received, blue is after 720 seconds of 1 kV argon sputter, black is after 2880 seconds of 1 kV argon sputter, and green is all previous exposure plus 30 seconds of 4 kV sputter. During the interrogation of the surface, sputtering reduced Zn to extremely low concentration and Mn was reduced as well. With the removal of these elements along with oxygen, the Cr, Fe, and Ni concentrations increased. The oxygen and carbon are never reduced to zero as the valleys on the surface of the powder particle are not exposed to the argon sputter. Although additional analysis is needed, all three powders analyzed displayed similar signatures regardless of cycle of reuse. The surface of the raw material is rich in Zn and Mn.

Although additional analysis is needed, all three powders analyzed displayed similar signatures regardless of cycle of reuse. The surface of the raw material is rich in Zn and Mn. Compared to a typical bulk stainless steel surface, Fe surface concentration is greater and Cr surface concentration is lower.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing and Technology, LLC under Contract DE-NA0002839 and Lawrence Livermore National Laboratory operated and managed by Lawrence Livermore National Laboratory , LLC under Contract DE-AC52-07NA27344356Virgin powder4 cycle powder, front overflow pre-sieveFigure 10: XPS surveys for various durations of sputterSEM was applied to characterize the surface of spatter particles produced by the metal AM process.

When material is ejected from the melt pool during active melting, the surface tension yields a spherical particle which, as it solidifies, generates a unique dispersion of elements. Elemental maps and spectrum identified oxide patches on the surface of these particles as silicon oxide (Figure 11 and Figure 12). These finding agree with work performed by Simonelli who noted Mn, Si, and Mg oxides on the surface of laser spatter particles [24].Figure 11: Elemental map of particleFigure 12: EDS of oxide (red) and base material (yellow)Absorptivity testing was performed on numerous samples and revealed noticeable differences in absorptivity relative to the sample collection location (Figure 13).

A linear fit was applied to the temperature-dependent absorptivity according to Equation 6.A(T) = α + βT(°C)Equation 6where α is the powder absorptivity at room temperature.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing and Technology, LLC under Contract DE-NA0002839 and Lawrence Livermore National Laboratory operated and managed by Lawrence Livermore National Laboratory , LLC under Contract DE-AC52-07NA27344357Alpha coefficientBeta coefficientFigure 13: Absorptivity coefficient scatterplots by collection location and powder cycleThe metal AM process certainly has an impact on the absorptivity of the powder. While there was some variation in the data, material collected from the exhaust duct and from the sieve screen (debris) typically yields higher values for absorptivity.

While there was some variation in the data, material collected from the exhaust duct and from the sieve screen (debris) typically yields higher values for absorptivity. As with particle size and composition, the majority of the material that is altered by the metal AM process is removed via sieving in between each reuse of the powder.Solidified SamplesInorganic analysis was performed on 57 solidified samples fabricated as part of this study. There was no consistent statistical trend in composition relative to powder reuse as displayed in Figure 14 for selected elements. Corrosion testing performed via copper sulfate found no indications of free iron on the selected samples.Figure 14: Solid article iCAP chemistry, selected test dataXRD analysis of samples displayed consistent lattice constant (3.615 ± 0.006 Ångström) and unit cell strain is high for metal AM items (0.99 ± 0.06).

The lattice constant was consistent with a typical austenitic structure with preferred orientation [25]. XRD revealed two prevalent crystal growth orientations. The preferential crystal growth direction was found to correspond to the (111) Miller index plane followed by the (200) Miller index plane.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing and Technology, LLC under Contract DE-NA0002839 and Lawrence Livermore National Laboratory operated and managed by Lawrence Livermore National Laboratory , LLC under Contract DE-AC52-07NA27344358Ninety-eight density samples were tested from twenty-eight builds.

Density had no consistent trend versus the cycle of reuse, implying that this characteristic is affected by slight variations in the AM process more than the powder reuse cycle.Figure 15: Density values determined via Archimedes in alcohol by powder cycle (average+/- 95% CI)One hundred eighteen tensile bars were tested in this study. There is no consistent trend in the mechanical strength and ductility relative to powder reuse for up to twelve cycles of reuse. A number of instances of low strength and ductility were observed throughout the study, but these were identified to be caused by various machine or process variations as opposed to being caused by the reuse of powder. The majority of these low strength results are associated with builds in which poor layer fusion was obtained due to either laser muting or laser focus drift. Laser muting is caused by condensate in the build chamber depositing on the lens cover and is compounded by poor gas flow due to filtration restriction.

Laser muting is caused by condensate in the build chamber depositing on the lens cover and is compounded by poor gas flow due to filtration restriction. Maintaining laser focal consistency with this equipment was a challenge, and in some instances, despite adequate chamber conditions and focal verification prior builds, was found to be the cause for lack of consolidation.Ultimate tensile strengthElongation % at breakFigure 16: As-AM selected tensile results©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing and Technology, LLC under Contract DE-NA0002839 and Lawrence Livermore National Laboratory operated and managed by Lawrence Livermore National Laboratory , LLC under Contract DE-AC52-07NA27344359Review of tensile fractures was performed on a limited number of samples, but provided items for future investigation.

A typical fracture is presented in Figure 17 with jagged edges around its circumference. Due to the surface finish of these parts, the failures propagated from numerous locations resulting in a non-uniform macro-cup and cone. Upon close inspection it was determined as ductile fracture due to the observance of microvoids or dimpled structure across the sample. These micro-cup and cone ductile fracture events are observed to be oriented relative to their position on the cross section, with fracture near the center of the sample equiaxed and oriented upwards, slowly changing to shear ductile rupture directionally oriented towards the center of the sample near the extents of the bar.39x5000xFigure 17: Typical fracture surfaceAn example of a non-typical tensile bar fracture that was investigated is presented inFigure 18. Due to poor layer adherence this failure appears to have a majority of cleavage rupture.

Due to poor layer adherence this failure appears to have a majority of cleavage rupture. Lack of solidification was supported by the observance of numerous regions of unmelted powder. Close observance appears to display the upper face of a weld pool, unchanged by the application of subsequent laser passes. This failure was caused by an extended build pause which was not appropriately restarted. During an extended pause, the components being fabricated cool to ambient temperature conditions. Upon restart, the build height is compensated for component thermal shrinkage such that the restarted layer is deposited directly on the last completed layer.

Upon restart, the build height is compensated for component thermal shrinkage such that the restarted layer is deposited directly on the last completed layer. This ensures weld penetration into the previously deposited layer, but was not performed according to the standard process in this instance.80x5000xFigure 18: Cleavage rupture example ©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing and Technology, LLC under Contract DE-NA0002839 and Lawrence Livermore National Laboratory operated and managed by Lawrence Livermore National Laboratory , LLC under Contract DE-AC52-07NA27344360One-hundred and fifteen samples were analyzed for microhardness using a 500g Knoop impression. Fifteen to eighteen measurements were taken per sample for a total of 1917 measurements.

Fifteen to eighteen measurements were taken per sample for a total of 1917 measurements. While there were some localized trends, no consistent statistical trend was observed in microhardness over twelve cycles of powder reuse (Figure 19). Metallography performed in parallel with microhardness observed no consistent trends in sample porosity (maintained below 0.15%) and observed consistent weld penetration depth and melt pool shape (Figure 20).Virgin12 cycleFigure 19: Microhardness data by cycle (ave ± 95% CI)Figure 20: Examples of sample metallographySurface finish was measured on twenty-two samples with approximately 11 linear inchesof readings taken on each sample. Surface finish of items produced was fairly consistent (382.6µin overall mean) regardless of powder cycle as displayed in Figure 21, with some exceptions. Builds occurring with two cycle powder had improved surface finish; attributed to tight gas flow monitoring that occurred over that period.

Builds occurring with two cycle powder had improved surface finish; attributed to tight gas flow monitoring that occurred over that period. A build using powder that had been reused eight times had poor surface finish; attributed to the metal AM machine becoming inoperable for an extended period of time over the course of that build.Figure 21: Surface roughness (R,a µin) by build end date©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing and Technology, LLC under Contract DE-NA0002839 and Lawrence Livermore National Laboratory operated and managed by Lawrence Livermore National Laboratory , LLC under Contract DE-AC52-07NA27344361Conclusions and Future WorkFor 316L stainless steel, the metal AM process does affect the powder used as observed via Rotap sieving, inorganic, SEM/EDS, and absorptivity testing performed.

However, the majority of the powder altered by the process is removed by the sieving process. In this study it was shown that 316L powder may be reused for up to twelve cycles in the laser powder bed fusion process with only minimal statistical variations in the powder size distribution and no consistent statistical variation in the properties reviewed by this study of the components fabricated by the metal AM process.Powder bed fusion has poor powder utilization at only 2-3%. Additionally, this process is still developing with significant operator intervention required to keep equipment processing and equipment issues resulting in build variation and intermittent equipment down-time.

In this study, the process variation associated with this emerging manufacturing technology had a more significant impact on results than did powder reuse.Future work should be focused on maturation of the metal AM process, reducing operator interaction and improving consistency of critical metal AM equipment settings such as laser energy density, gas flow, and powder layer thickness.

Additionally, efforts should include continued development of current powder and solid article analysis methods and the exploration of new methods to ensure repeatable results with suitable sensitivity relative to powder reuse.AcknowledgementsThe authors wish to acknowledge the support of numerous individuals in executing this work including Joseph Goedecke, Ida Sanchez, Joseph Hesting, Joe Morton, Joseph Vellella, Kevin Davis, Rella Greufe, Ann Choi, Emily Webster, Danielle Moore, David Smith, Ben Peterson, Mary Ann Godwin, Nathan Wyeth, Shawn Van Slyke, Ron Koch, Anthony Loveland, Gerald Woodburn, Brandon Volmerding, Justin Tannehill, Jason Gibbs, Jamie Ayers, Diana Goedecke, Jason Rogers, Andrew Cox, Gregory Hawkins, Michelle Maurer, Mary Urrutia, Bradley Teeter, Mark Sourek, Darryl Engelhardt, Michael Sutherlin, Paul Catlett, Karalyn Gibbs, Carol Lusby, Tamera Green, Joel Corkill, Jim Reilly, Michael Gatts, Karl Arnold, Phong Du, Otis Walton, Kaitlin Gill, and Kristen Roy.This work was funded and performed by Honeywell Federal Manufacturing and Technology under Contract DE-NA0002839 and Lawrence Livermore National Laboratory under Contract DE-AC52-07NA27344.

The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing and Technology, LLC under Contract DE-NA0002839 and Lawrence Livermore National Laboratory operated and managed by Lawrence Livermore National Laboratory , LLC under Contract DE-AC52-07NA27344362References[1]Seyda, V., Kaufmann, N., & Emmelmann, C. (2012). Investigation of aging processes of Ti-6Al-4V powder material in laser melting. Physics Procedia, 39, 425-431.[2] O'Leary, R., Setchi, R., Prickett, P., Hankins, G., & Jones, N. (2015).

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference3D BIOPRINTING OF SCAFFOLD STRUCTURE USING MICRO-EXTRUSION TECHNOLOGYJuan Xing, Xianli Luo, Juliana Bermudez, Matthew Moldthan, and Bingbing Li*Department of Manufacturing Systems Engineering and Management, California State University, Northridge, CA 91330-8332AbstractScaffold-based techniques are a vital assistance tool to support main structure andenhance the resolution of target structure. In this study, a custom-made micro-extrusion bioprinting system was built and utilized to fabricate different scaffold structures such as log-pile scaffold and two-ring scaffold. This approach showed tremendous potential because of its ability to produce microscale channels with almost any shape.

This approach showed tremendous potential because of its ability to produce microscale channels with almost any shape. We were able to fabricate these scaffolds by using a custom-made 3D bioprinter to print hydrogel solution, mostly composed of Pluronic F-127, then wash away hydrogen by phosphate buffer saline (PBS) after crosslinking of main structure. We were able to achieve the desired scaffold structure by feeding G-codes data into user interface (Pronterface) and then translating that model into a program that utilizes a customized programming language, which instructs the microfabrication printer nozzles to dispense the hydrogel at specific locations. This fundamental study will be used to print increasingly viable and complex tissue shapes with living cells.1. IntroductionOne of the foremost challenges facing scientists and researchers today in the field ofthree-dimensional (3D) bioprinting is the formation of viable and structurally sound tissues and organs.

The ever-growing demand for tissue and organ transplants has driven researchers to develop several different approaches to address this challenge. 3D bioprinting has opened the doors to new methods of printing, especially the methods used for tissue engineering and the production of artificial organs and tissues1. Having the capability to 3D print functional artificial tissues on demand is extremely beneficial for studying all aspects of these tissues and organs which may ultimately eliminate the need for donors2. Scaffold-based techniques provide an ideal support for cell adhesion, proliferation, distribution, and invasion3.

Scaffold-based techniques provide an ideal support for cell adhesion, proliferation, distribution, and invasion3. The structure of the scaffold can be modulated in order to re-create biological environments by arranging the scaffold’s spatial organizations that may be similar to in vivo tissues or organs4.Additionally, various formats and modifications to scaffold models, with respect to thechoice of biomaterials and physical conditions, have been implemented to optimize the growth of cancer cells on a scaffold5. It provides a way to mimic a natural tumor microenvironment which has become an approach to understanding cancer, while taking an insight into the potential motility of cancer cells in human bodies.A log-pile scaffolds is often used in cancer study for cell transmigration, where the spacing between the lines is based on the typical size of a cancer cell6.

The biggest challenge for 3D log-pile scaffold is to obtain high resolution, which is dependent of the width of the logs and the distance between the logs. For printing tissues and organs, the higher resolution of printed2311logs, the more sophisticated scaffold structure will be achieved. For the motility of cancer cells, they need micro-channels otherwise they tend to drop down to the bottom in the scaffold.In this study, a custom-made micro-extrusion bioprinting system was built and utilized for printing biological structures. This inexpensive method can be used to print desired shapes, and has ﬂexibility in the selection of the hydrogels used7. The bio-ink used is comprised of Pluronic F127 and glycerol, which was printed in a petri dish and maintained in room temperature before being used.

The bio-ink used is comprised of Pluronic F127 and glycerol, which was printed in a petri dish and maintained in room temperature before being used. Pluronic F127 with low biological toxicity and biocompatibility, has been a great candidate for cell encapsulation applications, either by itself or to specifically promote cell seeding and attachment in tissue scaffolds8. Glycerol, on the other hand, was used because of its well-known osmotic stabilizing properties9. As a result, Glycerol acts as lubricating agent to prevent clogging in micro-needles since glycerol can protect the bacteria against excessive stress caused by the encapsulation procedure. In order to obtain an optimal scaffold structure, we optimized parameters of the custom-made printing system, including needle diameter, interval between lines, extruding distances of piston for log and turning point, and extruding speed.

Hydrogel preparation, bioprinting system setup, results and discussion will be presented in the following sessions.2. Hydrogel PreparationAll solutions were prepared as a percent by mass. In a mason jar, Pluronic F-127 (P2443;Sigma-Aldrich) and glycerol (Sigma-Aldrich) were mixed in distilled water, resulting in concentration of Pluronic F-127 (33% w/v) and glycerol (10% v/v). The starchy mixture was stored at 4°C for 3 hours to obtain a liquid solution. The aqueous transparent solution was stirred for 1 hour, then was stored at 4°C before use. The liquid state of the solution provides convenience of application when it is being poured into a 20ml syringe. The syringe is then left at room temperature to start the formation of the non-occlusive gel on warming. The thermal gelation of Pluronic F-127 allows for a rapid printing due to its increase in contact time which produces a stable structure10.

The thermal gelation of Pluronic F-127 allows for a rapid printing due to its increase in contact time which produces a stable structure10. After leaving the material at room temperature for one hour, the solution is ready to be used for printing.Pluronic F127 aqueous solutions at 20 to 30% w/w has been known to having interesting characteristic of reverse thermal gelation11. Pluronic F-127 is at a liquid state when refrigerated at temperatures of 4-5°C, and gel upon warming to room temperature. The gelation is reversible upon cooling. The physical properties of this material makes it suitable to be used in micro- extrusion due to the ability to reverse thermal gelation. A hydration layer surrounds Pluronic F- 127 molecules at its aqueous state. As the temperature raises, the hydrophilic chains inside the copolymer become desolated which enables a breakage of the hydrogen bonds established between the solvent and these chains12. This phenomenon leads to gel formation.

This phenomenon leads to gel formation. This thermal reversible biopolymer gel has been widely used for pharmaceutical formulations and studies have demonstrated its high biocompatibility for biomedical applications, including grafts for wounds13 and drug delivery14-15.3. Bioprinting SystemThe design of the models was manually done by using a featured G-code sender with agraphical user interface (Pronterface). The software has a printer control section, a print simulation, a G-code sender, and a command prompt. The printer control section allows for auto-2312homing and moving of the axis. The print simulation area allows the user to see the design that will be printed based on the G-code made, and where in the platform. The G-code sender allows the user to directly send a G-code to the printer.

The G-code sender allows the user to directly send a G-code to the printer. The G-code, which provides existing standard protocols that can be easily modified, was manually written in order to control the intervals between the logs, the inner and outer diameter, as well as the supporting slopes. The G-code controls the tool pass and enables it to stop the process at each layer in order for solidification to occur.The biological structure processing in the study was implemented by using a micro-extrusion method executed on a custom built XYZ stages attached to a lid-screw syringe-based assembly that was custom designed in Figure 1. The syringe extruder was design for enabling precise hydrogel deposition. All the parts to make the custom syringe were 3D printed by a Stratasys uPrint SE 3D Printer. The syringe-based extruder uses an individual stepper motor independent of the XYZ axis platform to slowly move the plunger from a 20 ml syringe at a resolution of 500 micron.

The syringe-based extruder uses an individual stepper motor independent of the XYZ axis platform to slowly move the plunger from a 20 ml syringe at a resolution of 500 micron. Therefore, parameter modiﬁcations are needed to create a well-defined structure in terms of nozzle diameter, ﬁlament diameter, speed, flow, and G-code. A dispensing tip is used on the end of the syringe in order to control the volume of hydrogel being extruded. This method works by extruding a hydrogel material from the syringe to a petri dish. The petri dish was manually placed in the platform, while the extruder nozzle was positioned on top of the petri dish and the G-code of the desired design was sent to the printer. The printing process is set to slow in order to allow the solution to cool down and solidify on its own. When the printing process is concluded, the printed object is ready.Figure 1 Micro-extrusion Bioprinting System.4.

When the printing process is concluded, the printed object is ready.Figure 1 Micro-extrusion Bioprinting System.4. Fabrication of Two scaffold modes 4.1 Mode 1: Fabrication of log-pile scaffold2313Log-pile scaffolds are widely used in 3D bio-printing to fabricate artificial tissues and organs. Additionally, log-pile scaffolds enables to create a micro-environment so that it can be used to research the immigration path of cancer cells15. Therefore, in the both field of printing artificial tissues and organs, and exploring the metastatic path, log-pile scaffold plays a vital role.Log-pile scaffolds consist of several layers of 2D slab, with a previously formed logoriented perpendicularly. Since the extruding process was continuous, the width of an extruded log was always larger at the cross-sections due to the surface tension. Before printing log-pile scaffold with our micro-extrusion bioprinting system, we carried out a set of adjustment for 2D slab.

Before printing log-pile scaffold with our micro-extrusion bioprinting system, we carried out a set of adjustment for 2D slab. As long as we can get micron-scaled interval between logs, it will be easy to print sound log-pile scaffold with micro-channel.Adjustment in Setting Parameters for 2D slab In order to appropriately use the hydrogel solution on the printing systems there were various considerations that needed to be taken into account. We evaluated the hydrogel and bioprinting system based on five parameters that we believe would have an impact on the bioprinted structure. Specifically these are: distance from needle tip to plate surface (D, mm), moving distance of piston for one log (E1, mm), moving distance of piston for one turning short log (E2, mm), the speed of piston movement for one log (F1, mm/min), and the speed of piston movement for one turning short log (F2, mm/min).

All the data was taken in millimeters or millimeters per minute for dimensions and velocity respectively. A collection of these parameters for the evaluated trials is shown in Table 1.Table 1 Optimization of Printing ParameterTrial IDNeedle Diameter (mm)Interval between logs (mm)DE1E2F1F210.7210.10.140040020.51.510.070.07 40040030.51.20.80.050.05 40040040.51.20.50.050.05 40040050.51.20.50.050.03 40040060.51.20.50.030.03 40040070.51.00.50.020.02 40030080.51.00.30.020.02 4003002314abcdFigure 2. Log-Pile Structure. (a) The structure (30mm*15mm) shows the optimal 2D slab with 0.5mm of log diameter and 0.5mm of interval. (b) The 2D slab was extended into a square slab (30mm*30mm). (d) A two-layer log-pile scaffold where second layer was rotated by 90°.

(b) The 2D slab was extended into a square slab (30mm*30mm). (d) A two-layer log-pile scaffold where second layer was rotated by 90°. (d) A six-layer log-pile scaffold (30mm*30mm*2mm) was built by repeating C for three times.After the log-pile scaffold with 0.5mm of log diameter and 0.5mm of interval was successfully printed, we continued to reduce the micro-channels. In order to fabricate the log-pile scaffold with 0.5mm of log diameter and 0.3mm of interval, we adjusted the parameters. Based on several trials, the optimal parameters in settings were 0.3mm of the initial distance from needle tip to plate surface, 0.5mm of Interval between logs, 0.019mm of E1, 0.019mm of E2, 400mm/min of F1, and 300 mm/min of F2. The log-pile scaffold with smaller micro-channels was printed out with the parameters, was showed in Figure 3.b Figure 3. Log-Pile Structure. (a) The square structure (30.4mm*30.4mm) shows the optimal 2D slab with 0.5mm of log diameter and 0.3mm of interval.

Log-Pile Structure. (a) The square structure (30.4mm*30.4mm) shows the optimal 2D slab with 0.5mm of log diameter and 0.3mm of interval. (b) A two-layer log-pile scaffold where second layer was rotated by 90°. (c) A six-layer log-pile scaffold (30.4mm*30.4mm*2mm) was built by repeating B for three times.ac4.2 Mode 2: Fabrication of ring scaffoldsVascular-like constructs have been successfully printed in many tissue engineering labs, and it is a common way to fabricate it in log-pile scaffold16. However, we have not seen any labs utilize ring scaffold. Saving scaffold material is one of the significant advantages for ring scaffold.2315Figure 4. Ring Scaffolds. (a-1) The first layer of scaffold (a-2). The pattern was rotated by 22.5 degrees every other layer. (b-1) The first layer of scaffold (a-2). The pattern was rotated by 45 degrees every other layer.There were six layers in both scaffold a-2 and b-2.We created two types of ring scaffold a-2, b-2 showed in Figure 4.

The pattern was rotated by 45 degrees every other layer.There were six layers in both scaffold a-2 and b-2.We created two types of ring scaffold a-2, b-2 showed in Figure 4. These two have the same diameter of outer circle (20mm), same diameter of inner circle (12mm), same thickness (4mm), and the same height (2mm). The only difference is the number of crossed lines. Scaffold a-2 has twice number of zigzag lines as b-2, which results in different traits and functionalities. Scaffold a-2 can provide more support for main structure while Scaffold b-2 can create more space for the material of main structure.5. DiscussionIt is the crossed lines that enhances the stability of scaffold and creating support for mainstructure. For mode 1, log-pile scaffold, the 2D structure in Table 1 shows the process we reduced the log’s diameter and intervals between two logs.

For mode 1, log-pile scaffold, the 2D structure in Table 1 shows the process we reduced the log’s diameter and intervals between two logs. The 2D structure in Figure 3 indicates the smallest interval between logs (300 microns) we can get so far with our micro-extrusion bioprinter. Based on one layer of paralleled logs we printed 6 layers logs-piled structure shown in Figure 2 and Figure 3. The even layers and odd layers were interleaved by 90◦, forming a polyporous structure. For mode 2, ring scaffold, we created two different rings via designing different first layer and rotating the even layers by different degrees, showed in Figure 4. In mode 1 we used thinner needle with diameter of 0.5mm as there were so many thin logs crossed that the scaffold was stable enough.

In mode 1 we used thinner needle with diameter of 0.5mm as there were so many thin logs crossed that the scaffold was stable enough. However thicker needle with diameter of 0.7mm was used in mode 2 since we need build thicker inner and outer wall for ring scaffold or it was likely to collapse.In the parameter settings, D, E1, E2 were three critical parameters that needed to beiteratively adjusted. If D, distance from needle tip to plate surface, was too high, the part near turning section would become messy. If D was too low, the logs would become too thick, that was likely to reduce the interval space. E1 and E2 controlled the amount of extruded material for the log between two specific points, but the value of E1 and E2 were not proportional to the length of corresponding log due to the inertia effect of extrusion.23166.

Conclusion and Future WorkIn this study we optimized parameters of the custom-made printing system, including needle diameter, interval between lines, extruding distances of piston for log and turning point, and extruding speed. This fundamental study will be used to print increasingly viable and complex tissue shapes with living cells.Future work may include the following: testing lower concentrations of bacterial cellulosic exopolysaccharide gel on this micro-extrusion bioprinter to test the compatibility; improve the XYZ stages and dispensing nozzle parameters and increase accuracy and transparency of alginate models; inclusion of cells in hydrogel to test cell viability.7. AcknowledgementsThis research work was supported by the NIH NIGMS BUILD (Building Infrastructure Leading to Diversity) PODER (Promoting Opportunities for Diversity in Education and Research) Faculty Scholar Academy and Mentor Training (1RL5GM118975), Equipment (8UL1GM118976) and Student (8TL4GM118977).

 inkjet printing of conductive structures using in-situ IR sinteringE. Saleh*, J. Vaithilingam, C. Tuck, R. Wildman, I. Ashcroft, R. Hague, P. DickensUniversity of Nottingham, Additive Manufacturing and 3D Printing Research Group, University of Nottingham, Nottingham, UK, NG7 2RDAbstractIn this study we investigate the inkjet printing of a silver nanoparticle ink and the optimization of IR sintering conditions to form 3D inkjet-printed conductive structures. The understanding of the interaction between the silver layers and the sintering conditions are key elements to successfully build conductive tracks in 3D. The drop size of conductive ink on glass substrates as well as on sintered conductive film was measured to optimize the printing resolution. The resistivity of the sintered deposition was studied in a planar X-Y direction as well as in a vertical Z direction to analyze the effects of stacking hundreds of silver layers in different deposition orientations.

Using the results of the optimized printing and sintering conditions, conductive tracks were demonstrated forming simple 3D inkjet-printed structures powering electronic components.Introduction3D printing of electronic components and circuitries are key target areas of additive manufacturing (AM) technologies. Taking electronics outside the traditional two-sided or limited layers printed circuit boards (PCB) into the 3rd dimension will enable various new electronic applications. Similar to 3D integrated circuits (3DIC), compact and flexible electronics are amongst those applications that 3D technologies could enable, where conductive tracks can move in a volumetric manner rather than a 2D surface allowing more freedom of track routing [1-3]. Antenna design is another area that has benefited from 3D technologies where conductive tracks printed on 3D surfaces showed better performance and higher efficiency antenna than conventional fabrication methods[4, 5].

Most common form of conductive inkjet inks are precursors of silver nanoparticles dispersed in solvents [6, 7]. A challenge to 3D-print conductive inks is the sintering process which is usually a heating process to evaporate the solvent of the conductive ink and sinter the nanoparticles to form a percolated path[6]. In this study we report using infrared (IR) heating to sinter multilayer tracks of inkjet-printed silver ink. IR sintering has been reported in roll-to-roll processes where the inkjet-printed patterns are sintered using high power IR lamps to form 2D conductive circuits [8-11].ExperimentalIn order to understand the printing conditions the silver ink was jetted at different substrates to optimize the droplet size and the printing resolution. Upon inkjet printing, the IR lamp was configured to sinter the deposition when each layer of deposition is complete. Tracks containing layers from 1 to 20 were printed and IR sintered using these conditions.

Tracks containing layers from 1 to 20 were printed and IR sintered using these conditions. The resistance of thetracks was measured as the number of layers increase. Finally 300 layers of an array of single pixels were printed to study the effect of printing large number of layers on the overall resistance in the z direction and compare it to the x-y planer patterns.Materials: A commercial silver nanoparticle conductive ink from Advanced Nano Product (ANP) – product number: DGP 40LT-15C - was used in this study. The ink contains 38.5 wt% of <50 nm silver nanoparticles dispersed in triethylene glycol monomethyl ether (TGME). Glass slides used as substrates were soda lime glass slides from Cole-Parmer Instrument Co. Ltd.Apparatus: The ANP silver ink was inkjet-printed from a Spectra SE128 print-head installed in a bespoke printer from PixDro (commercial name Toucan).

Ltd.Apparatus: The ANP silver ink was inkjet-printed from a Spectra SE128 print-head installed in a bespoke printer from PixDro (commercial name Toucan). An IR lamp from Heraeus (Heraeus 4114) was installed 800 mm away from the print-head in the direction of printing, which occurs by moving the substrate under the IR lamp for a number of passes as shown in figure 1.Figure 1, Toucan inkjet printer with IR sintering unit.The jetting conditions of the ANP ink were 90 V jetting pulses each is 12 µs (2 µs ramp, 8 µs dwell time and 2 µs falling time). The ink was jetted at 30 ºC to tune the viscosity, and the ink pressure on the print-head was maintained at -20 mBar.

The ink was jetted at 30 ºC to tune the viscosity, and the ink pressure on the print-head was maintained at -20 mBar. The height of the print-head from the substrate was 1 mm and the height of the IR lamp was kept at 7 mm away from the substrate which corresponds to a temperature of 290 ºC measured at the surface of a glass substrate.ResultsDrop size and printing resolution: The size of the droplet size at the substrate is determined by many factors, one is the substrate material. In 3D inkjet printing, using a glass substrate as an example, the droplet size on the glass substrate would change in the second layer after the first layer of the ink was formed if the materials have different surface energies. To evaluate the printing resolution the droplet size of the ANP ink was measured on soda lime glass slide and also on sintered silver ink.

To evaluate the printing resolution the droplet size of the ANP ink was measured on soda lime glass slide and also on sintered silver ink. Figure 2 shows a comparison of the droplet size on the two materials.Figure 2, Droplet size of the ANP conductive ink on a) soda lime glass substrate. b) Sintered silver substrate.The average droplet size of the ANP on glass was 118 µm whereas on silver it was 145 µm. Assuming droplets overlap of 70% the resolutions on glass and silver substrates are 308 dpi and 250 dpi, respectively.Electrical characterization: To measure the electrical resistivity of the sintered ink tracks of 7 by 0.5 mm were printed and sintered using the printing conditions described in the experimental section.

Square pads were also printed at the end of the tracks to act as measurement points as shown in figure 3.Figure 3, Multi-layer tracks of IR sintered silver ink printed on glass substrate (5mm scale bar).The resistance of each track was measured using Hameg LCR high precision meter to evaluate the development of the resistance as the number of layers increase. Figure 4 shows the average resistance of the tracks taken from multiple samples; each track was sintered by applying 10 passes of IR over the deposition which corresponds to 60 seconds of IR exposure.Figure 4, The resistance of the tracks as it changes with the number of layers.A number of observations can be made on figure 4. The change in resistance with the increase of layers is nonlinear although in theory the increase is geometry dependent so assuming the volume of each layer is the same the resistance trend should have been linear.

This nonlinear trend can be due to different deposition volumes per layer as the layers increase; however the most likely reason of this nonlinearity can be due to the presence of organic residue between the layers. The organic materials between the layers usually come from any additives in the ink other than the solvent (surfactants and viscosity modifiers) which don’t evaporate during the sintering process. To measure the resistance of the sintered ink vertically 300 layers of single pixels array were printed and sintered on brass substrate using the same printing and sintering conditions as the planer tracks. The resistance of the tracks was measured between the brass substrate and the tip of the silver as shown in figure 5.Figure 5, The configuration of measuring the resistance of 300 sintered silver layers vertically.The measured resistance of the 300 layers was 21±4 mΩ. The diameter of each pixel was around 150 µm and the height was around 200 µm.

The diameter of each pixel was around 150 µm and the height was around 200 µm. Comparing that to the estimated resistivity for less number of layers (1-20 layers), the overall resistivity increases when the number of layers increase. This confirms the influence of the interlayering organic materials acting as small resistors between the layers. The conclusion of the presence of organic interlayers needs further study to include an accurate surface profile of the tracks to measure the resistivity quantitatively and compare it to higher aspect ratio silver pillars in order to draw a further rigorous conclusion and a mathematical model of the resistivity as it changes with the number of layers.ConclusionPrinting and sintering multilayer conductive ink was achieved using an inkjet print-head with IR post-processing. The droplet size on glass substrate and silver film was measured to evaluate the influence of the substrate and determine the optimal printing resolution.

The droplet size on glass substrate and silver film was measured to evaluate the influence of the substrate and determine the optimal printing resolution. Multiple layers of silver tracks were printed and sintered for each layer at the same conditions. The resistance of tracks was measured and showed a nonlinear behavior. A hypothesis of this nonlinearity as the number of layers increase was introduced. Organic interlayers from various additives in the ink add a fine resistance between the silver layers increasing the overall resistivity of the structure.

Organic interlayers from various additives in the ink add a fine resistance between the silver layers increasing the overall resistivity of the structure. This conclusion needs further investigation to quantitatively measure the resistivity of the sintered tracks in addition to the vertical resistance that was introduced in this study.AcknowledgmentsThe authors gratefully acknowledge funding support from EPSRC (Engineering and PhysicalSciences Research Council) under the Jetting Electronic Tracks JET project, and the EPSRC Centre for Innovative Manufacturing in Additive Manufacturing (EP/I033335/2) at the University of Nottingham.References1.2.3.4.5.6.7.8.Davis, W.R., et al., Demystifying 3D ICs: the pros and cons of going vertical. Design & Test of Computers, IEEE, 2005. 22(6): p. 498-510. Xiangyu, D. and X. Yuan. System-level cost analysis and design exploration for three- dimensional integrated circuits (3D ICs). in Design Automation Conference, 2009.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference3D INKJETTING DROPLET FORMATION OF BACTERIAL CELLULOSIC EXOPOLYSACCHARIDE GELDaniel Aguiar*, Amanda Albuquerque†, and Bingbing Li*Department of Manufacturing Systems Engineering & Management, California State University Northridge, Northridge, CA 91330 †Universidade Federal de Pernambuco - UFPE, Pernambuco, BrazilAbstractOn-demand 3D printing of scaffolds and cell-laden structures has shown promising results that can significantly impact human welfare. The objective is to fully understand the behavior of bacterial cellulosic exopolysaccharide gel (BCEG) as a new bioink with low toxicity and high biocompatibility for regenerative medicine.

The objective is to fully understand the behavior of bacterial cellulosic exopolysaccharide gel (BCEG) as a new bioink with low toxicity and high biocompatibility for regenerative medicine. Its possible application is to construct scaffolds that can be used for several biomedical applications, especially tissue engineering and treatment of critical bone defects. By using a MicroFab inkjet micro dispenser, BCEG was dispersed to create drops on demand that can be used to fabricate scaffolds. In order to fully understand the material’s behavior and droplet formation, we analyzed the physical and mechanical properties of the BCEG in different concentrations (0.1% 0.5% and 1%) and characterized it by its macroscopy, microscopy, rheology and particle size distribution.1. IntroductionWithin the past few years, the manufacturing world has undergone several changes, including the creation of new materials and new manufacturing techniques.

IntroductionWithin the past few years, the manufacturing world has undergone several changes, including the creation of new materials and new manufacturing techniques. Among them, we can highlight the field of additive manufacturing, also known as 3D printing, which has advanced continuously through the past few decades1. 3D printing is widely applied in various industries, including the biotechnology field, and scientists are constantly improving it. 3D bioprinting in particular has shown promising results and it is being studied extensively, in order to further refine its application and biomimicry accuracy.Currently, there are few 3D bioprinting techniques in practice and each of them have their advantages and disadvantages. Among them all, there are three types that have been used extensively in studies: micro extrusion, inkjetting and laser-assisted2. Inkjetting is one of the preferred methods for dispensing cells because of its high cell viability.

Inkjetting is one of the preferred methods for dispensing cells because of its high cell viability. This is due to the lower shear stress caused during the dispensing of bioinks, when compared to other methods such as micro extrusion1.The objective of this study is to analyze and understand the droplet formation process when dispensing the cellulosic exopolysaccharide gel (BCEG) through a micro nozzle dispensing system, using one of the inkjet approaches called Drop on Demand (DOD). Ultimately, the motivation of this experiment is to decide whether it is a viable method for scaffold construction, which can be used in several biomedical applications. These uses include tissue engineering and treatment of critical bone defects, as bone is the second most transplanted tissue worldwide.3This is a promising technique for in situ fabrication of scaffolds that can be customized to patient needs and specific geometry.

The goal is to create a highly compatible scaffold by using the cellulosic exopolysaccharide gel. This would ultimately result in reduced surgery time and decreased immune rejection, while presenting a viable solution for the current problem of organ donor shortage. This use of novel renewable material is very promising and manufacturing1687technique approaches shall be studied in order to achieve an effective, repeatable and scalable fabrication.2. Materials and Methods 2.1 Hydrogel FabricationThe proprietary Bacterial Cellulosic gel, POLYGEL®, is a hydrogel produced by POLISA, a startup originated at the Biopolimero de Cana-de-Açúcar Group at the Universidade Federal Rural de Pernambuco. This gel is a natural product composed of bacterial cellulosic exopolysaccharide obtained from sugarcane molasses by Zoogloea sp4. This polymer was refined in PA grade and was presented as a hydrogel in three concentrations: 0.1%, 0.5% and 1%. w/v.

This polymer was refined in PA grade and was presented as a hydrogel in three concentrations: 0.1%, 0.5% and 1%. w/v. The solutions were solely composed of dry bacterial cellulose and deionized surgical grade water, and sterilized by gamma-ray irradiation under ANVISA standards, for use in this study.This bacterial cellulosic exopolysaccharide biopolymer has been used for different biologic applications and studies have demonstrated its high biocompatibility for biomedical applications, including grafts for wounds and bone defect repair in animal studies5, 6 and in humans7, 8, 9. The biodegradability, sustainability and wide availability of its raw materials are also very interesting properties of this novelty material. When compared to materials already available on the market, such as ePTFE, the BCEG is a cheaper yet more effective material to construct films and grafts10.

When compared to materials already available on the market, such as ePTFE, the BCEG is a cheaper yet more effective material to construct films and grafts10. Due to the fact that hydrogels have high water content, they are ideal for absorbing high levels of nutrients and oxygen11. Its high biocompatibility and bio integration make hydrogels a strong material candidate that can be applied to a wide range of medical applications.2.2 Experiment SetupThe inkjet printing process in this study was implemented using a MicroFab PH-46(MicroFab, Plano, TX) micro dispensing subsystem. The bioink was ejected using a MicroFab 80 μm micro nozzle dispenser. The DOD pulse was controlled using a MicroFab Jet Driver, and a MicroFab pneumatic controller was used to adjust the backpressure of the fluid reservoir, to obtain an ideal meniscus for good droplet formation. An imaging system camera STC- MB33USB (SENTech America, Carrolton, TX) was used to observe the droplet formation process.

An imaging system camera STC- MB33USB (SENTech America, Carrolton, TX) was used to observe the droplet formation process. The setup schematic is shown in Figure 1. The parameters were adjusted for each of the gel concentrations, in order to obtain the droplets by increasing or reducing the back pressure and pulse. Initial arbitrary values were used, in order to set a range for the polymeric gel used in the study, then were fine-tuned.2.3 Hydrogel CharacterizationMacroscopic observations were made and the hydrogel’s visual, olfactory and tactilecharacteristics were recorded. Microscopic analysis was performed by Scanning Electron Microscopy Zeiss Supra 40VP (Zeiss, Oberkochen, Germany) and utilized to obtain images of the dried polymer samples’ morphology and organization. Optical Microscopy (Olympus, Tokyo, Japan) imaging was also performed, in order to observe the arrangement of the hydrated fibers in the hydrogel.

Optical Microscopy (Olympus, Tokyo, Japan) imaging was also performed, in order to observe the arrangement of the hydrated fibers in the hydrogel. A Microtac Zeta Potential analysis (MicrotracBel, Osaka, Japan) was performed and the particle size distribution was observed for the biopolymer solution concentrations. Rheology of the hydrogel was characterized using a TA AR200 Rheometer (TA Instruments, New Castle, DE), using a 20mm parallel plate geometry with a measuring gap of 1 mm, and experimental conditions were kept constant. Temperature was kept constant at 25 degrees Celsius, as viscosity of the hydrogel can vary depending on temperature12.1688Figure 1 Microfab dispensing experiment setup.3. Results3.1 Macroscopic ExaminationDuring the visual macroscopic examination of the polymer at room temperature, it presented a homogeneous, semi-transparent, viscous aspect.

Results3.1 Macroscopic ExaminationDuring the visual macroscopic examination of the polymer at room temperature, it presented a homogeneous, semi-transparent, viscous aspect. Its olfactory properties presented no smell and the gel was soft to the touch, with a gel-like texture.3.2 Microscopy Analysis 3.2.1 OpticalWidefield microscopy was conducted and the hydrated fibers in the solution were observed. They presented a uniform morphology of elongated fibers that were clumped together in clusters. On Figure 2, a fibril cluster is observed in a 1% BCEG sample.Figure 2 Fibril cluster of the Bacterial Cellulosic Exopolysaccharide at 1% concentration16893.2.2 SEMThe BCEG samples were dried for 24 hours over silicon wafers on a heat plate set to 45 degrees Celsius. It was then gold coated and taken to SEM imaging in Figure 3. The images show the morphology and arrangement of the polymer fibrils when dried.

It was then gold coated and taken to SEM imaging in Figure 3. The images show the morphology and arrangement of the polymer fibrils when dried. Their size reduced due to shrinkage during the drying process where the water content was evaporated.Figure 3 SEM imaging of dry samples of the polymer showing the morphology and arrangement of its fibers.3.3 Zeta PotentialThe zeta potential test shows that the particle size range (0.05 - 6.54 um) was the same in all samples as the only variable changed was the concentration of the polymer.

The particle size distribution is somewhat uniform in each polymer concentration in the hydrogel form.Figure 4 Zeta potential particle size distribution in 0.1%, 0.5% and 1% concentrations of the biopolymer3.4 RheologyThe Viscosity as a function of shear rate for the BCEG in 0.1%, 0.5% and 1%concentrations recorded using the rheometer in Figure 5.1690Figure 5 Viscosity and shear ratio of the BCEG in 0.1%, 0.5% and 1% concentrations3.5 Ideal DispensingWhen the drops that are formed fall in a controlled fashion, as it is observed for the 0.1%concentration BCEG gel in Figure 6.Figure 6 - An ideal droplet sequence formation of BCEG in 0.1% concentration3.6. Non-Ideal DispensingNon-ideal droplet formation is observed, with the incorrect droplet dispensing formationresulting in inaccurate geometry, in Figure 7.1691Figure 7 Non-ideal droplet formation in 0.5% BCEG concentration4.

DiscussionAs expected, the gel presented higher viscosity proportional to the concentration of exopolysaccharide of the polymer. Figure 5 illustrates the viscosity of the gel and its shear thinning properties for concentrations of 0.1%, 0.5% and 1%, respectively. The polymer with 0.1% concentration behaves like water, with very low density. The viscosity will affect the how the gel is dispersed through the micro dispensing nozzle.When ideally formed, the drops will fall in a controlled fashion. Each drop’s size and rate of falling are both uniform, and these properties can be controlled by changing the dispensing parameter, which is a combination of back pressure, microcontroller and pulse control. In Figure 6, the sequential formation of uniform droplets is observed for the 0.1% concentration BCEG gel, which has the lowest viscosity among the samples used in this study.

In Figure 6, the sequential formation of uniform droplets is observed for the 0.1% concentration BCEG gel, which has the lowest viscosity among the samples used in this study. The higher the concentration of the gel, the more difficult it becomes to achieve an ideal droplet formation.As presented in another study1 non-ideal droplet formation is frequently observed. With the incorrect droplet dispensing formation, the accuracy and precision of the drop fall is compromised, so the drops will not stack or align with the previous droplet, resulting in inaccurate geometry. During the dispensing process, the polymer’s micro fibrils are partially ejected through the nozzle attaching to the tip, which disturbs the trajectory and formation of the drops. This was clearly observed when dispensing the hydrogel in 1% concentration as the accumulation of particles in the nozzle is directly proportional to the viscosity of the polymer solutions.

This was clearly observed when dispensing the hydrogel in 1% concentration as the accumulation of particles in the nozzle is directly proportional to the viscosity of the polymer solutions. The viscosity and the fibrous particle size, as well as the accumulation in the nozzle tip, are also responsible for the disturbance and can be observed in Figure 7. Wetting of the nozzle tip is another problem that commonly occurs while inkjet printing, since the back pressure must be carefully adjusted before dispensing, contributing to this issue.Based on the collected data, the droplet formation is relatable to its rheological properties; the higher the viscosity of the hydrogel, the more difficult it is to dispense it3. The hydrogel in the concentration of 0.5% presented very inconsistent droplet formation and the ideal formation was not observed in the concentration of 1% either.

The hydrogel in the concentration of 0.5% presented very inconsistent droplet formation and the ideal formation was not observed in the concentration of 1% either. The polymer micro fibrils tended to accumulate in1692the tip of the dispensing nozzle, making it more difficult to dispense the material. In addition to that, the fibrils passing through the nozzle changed the shape of the droplets. The SEM and optical imaging confirms the assumption that the fibrils are elongated and tangled. Analyzing the viscosity of the three concentrations shows a much higher value for the 0.5% and 1% concentrations, when compared to the 0.1%. This information alone can justify the DOD behavior and uniform formation when dispensing the polymer in lower concentrations which has a much lower viscosity and fluidity.5.

This information alone can justify the DOD behavior and uniform formation when dispensing the polymer in lower concentrations which has a much lower viscosity and fluidity.5. Conclusion and Future WorkKnowing the rheological properties and inkjetting behavior of the BCEG polymer was crucial in order to fully understand and design an effective dispensing mechanism to fabricate 3D structures for medical applications. With a uniform and on-demand dispensing system, a repeatable and scalable automated system could be designed and constructed to fabricate scaffolds. Unfortunately, the inkjet Drop on Demand system might not be the best alternative to disperse the biopolymer gel, due to its high viscosity variation. The size of the fibrous microstructure and the difficulty of configuring the inkjet parameters for each different concentration also complicated the process in which ideal dispensing was only reliably achieved when dispensing the polymer at 0.1%.

A pneumatic extrusion method shall be tested and validated, due to its better handling of gels in several viscosities and the fact that it is easier to control the dispensing parameters and exchange dispensing tips. A study of the relation between the biopolymer concentration and tissue regeneration is in progress and will be published later this year. That will be crucial to determine the ideal gel concentration for dispersion, based on cell viability, organism assimilation and regeneration factors.6. AcknowledgementsThe authors gratefully acknowledge the CSUN NIH BUILD PODER Equipment Competition Grant: 8UL1GM118976. The Biopolimero de Cana-de-Açúcar group and Professor Dr. Jose Lamartine Aguiar at the Universidade Federal de Pernambuco, Brazil. Professor Dr. Xiaochum Li and Mr. Injoo Hwang at the Scifacturing Lab at UCLA for the SEM and Optical imaging. Professor Dr. Pirouz Kavehpour and Ms.

go Ederer, Rainer Hochsmann, JUrgen Machan, FOB, Technical University of MunichAbstractMany of the currently available RP-Systems are suitable for building design models of arbitrarily shaped parts. However, most of these RP processes use sophisticated and expensive equipment which is not well suited for an office environment. In this paper we present a method and an experimental device for building design models by a modified 3D print process using plastic powder and a photopolymeric binder.Introductionrecent time RP-technologies are increasingly viewed under the perspective of their ability to create functional prototypes or tools. On the other hand, potential which arose from the ability to produce cost-effective design models is still not fully exploited. The advantages of such parts are clearly visible: one can avoid construction errors, solve assembly and production there are some special requirements to an problems, or compare design alternatives.

The advantages of such parts are clearly visible: one can avoid construction errors, solve assembly and production there are some special requirements to an problems, or compare design alternatives. However, "office-compatible" process which are not completely met at this moment by commercially available RP systems. One very important requirement for such a device is that it should be as easy to use for a designer as his plotter for graphic output. Therefore, it has to fit smoothly into the working environment of a design office. That means that its dimensions should be comparable to other office devices, that there are no special interfaces or media, and that waste materials and all other emissions are neglectable. Such a device should be easy to use and operate under the control of a CAD workstation. Cost of ownership has to be in the range of other office equipment.

Such a device should be easy to use and operate under the control of a CAD workstation. Cost of ownership has to be in the range of other office equipment. sufficient rigidity, The design models should meet the typical requirements for such objects, mid-range accuracy, good surface quality which eliminates the need for finishing work, and - if possible - compatibility with follow-up processes. An additional requirement would be the ability to recycle design models which are no longeruse.Experimental set-upTo test this hypothesis, an experimental setup as described below was designed and built at The above mentioned criteria were applied to the selection of materials for 3D printing. We used powders of polyamide and polystyrene as fillers, and UV-curable resins like epoxides and vinyl ether as binders.

We used powders of polyamide and polystyrene as fillers, and UV-curable resins like epoxides and vinyl ether as binders. The main components of the experimental setup are:e • ..fIlan x-y-plotter which is equipped with an ink-jet printhead for binder application, a building platform which moves downward from the starting level in small steps, a recoating mechanism for the spreading of new powder layers, an UV illumination unit above the platform.130",A " "modelframepiston...,. .....'A .......'A...h process is performed by repetition of a sequence of our process, thet"\11"\rl11411"T by by a UV lampdroplet-generator. completion ofhad to is a commercially available printhead employs piezoelectric actuators to move one of chamber by other side. diameter isto an electric signal1).

completion ofhad to is a commercially available printhead employs piezoelectric actuators to move one of chamber by other side. diameter isto an electric signal1). to an ejectionon one sidel"n~lrnlr'\""'1'" measures about 2 x 2 x 0.2 mm3, theMost ofliquid path is made of glass, which gives the printheadstability weWe take advantage ofetc., as in continuous jet devices,principle, drop131a)b)3Because our printing system was designed and optimized for ink, any other usable pumping liquid must be similar to ink with respect to three material properties: • dynamic viscosity:1...10 mPas (milli-Pascal-seconds) 72 mN/m (at 20°C) ca. 1000 kg/mJsurface tension:density:c)d)Fig. 2: operating principle of a piezo driven micropump5The dynamic viscosity of preferable resins for our 3D printing process is about 80 mPas at room temperature, which is far too high for the selected printhead.

Fortunately, the viscosity drops rapidly with rising temperature; at 60°C we achieve 10 mPas, which is a reasonable value for ink-jet printheads. The modified printhead is equipped with a resistive heater, a temperature sensor, and a temperature control circuit.Thorough investigation of the driving pulse for the piezoelectric actuator revealed the need for a specially adapted timing for ejection of resins. A freely programmable high-voltage pulse generator gave valuable help in these experiments. Following drop ejection characteristics have been achieved so far:• • •avg. drop mass: avg. drop diameter: ejection frequency: drop velocity:ca. 125 ng (nanograms), ca. 60 microns, 1000 Hz, ca. 2 m/s.Drop size and ejection frequency are key values for the overall performance of the RP system, because they define the model building speed via the deliverable amount of binder per unit of time. The table below shows this context.

The table below shows this context. These calculations are based on the assumption that 30 percent of the model volume are binder resin. The given dosing time refers to the building of a cube with 100mm edge length.frequency [Hz]number of nozzlesdosing time [h]100017371000501550005035000.1001,5By multiplying the operating frequency and usmg multi-nozzle arrangements, performance will grow dramatically.the system132Similar to some SLS (selective laser sintering) processes, powder coating is performed by a blade mechanism, which covers every cured layer with a precisely defined new powder layer from the reservoir. After application of the binder pattern for a specific layer, the entire powder-binder surface will be exposed to ultraviolet radiation; thereby the binder resin will be cured rapidly. The employed photopolymers in our experimental setup need a wavelength of about 350...365 nanometers and a specific curing energy of 70...

The employed photopolymers in our experimental setup need a wavelength of about 350...365 nanometers and a specific curing energy of 70... 100 millijoules per square centimeter. A reflector-equipped, commercially available UV lamp is mounted at some distance on top of the setup and does a very good job. The lamp delivers about 20 W power in the desired wavelength range; the exposure time for a 200 mm x 200 mm area is only 2 seconds. This time frame can be used for printhead cleaning after completion of each layer.ConclusionWe could demonstrate a 3D printing process and a 3D printing device which is made up from very simple components and is suitable for creating arbitrarily shaped plastic design models. The total cost of the device (without computer) is well below DM 5000,- (US-$ 3500.- ). Despite the selection of a rather expensive photopolymer, running cost is also very low, because no laser system is needed.

Despite the selection of a rather expensive photopolymer, running cost is also very low, because no laser system is needed. Low price and compact dimensions of the device make it ideally suited for desk-top use.AcknowledgmentsThe authors would like to thank Dr. Heiner Bayer and Walter Fischer of Siemens AG Corporate Research and Development Base Technologies Materials and Recycling for their tireless support.References[1]Heinzl, 1., C.H. Hertz: "Ink-Jet Printing", Advances in Electronics and Electron Physics, Vol. 65, PP91-171, Academic Press 1985[2]Berchtold, A.: "Optimization of a Drop-on-Demand Print Head with Planar Piezoceramic Transducers", SID Symposium, Las Vegas 1990[3]Cima, M. 1., Sachs E., "Three Dimensional Printing: Form, Materials and Performance" Solid Freeform Fabrication Symposium, Austin 1991[4]Sachs, E., Brancazio, D.: "Three-dimensional printing techniques", International Patent WO 93/25336 June 4, 1993133.

 Printable Wax-Silicone Actuators Jeffrey Ian Lipton*, Sarah Angle, Hod Lipson Cornell Creative Machines Lab, Cornell University *Corresponding author: jil26@cornell.eduAbstract: The Solid Freeform Fabrication of actuators has been an area of active development. So far only weak polymer actuators, or small displacement piezoelectric, and pneumatic actuators have been produced. We developed a novel material platform of silicone and wax which can be used to make soft actuators that are thermally activated. The material is made by mechanically mixing liquid silicone and liquid paraffin wax and cooled to create a suspension of wax particles suspended in a silicone liquid. The resulting material expands by up to 6% of volume when heated above the wax melting temperature.Keywords: Wax, Actuator, Robocasting1. Introduction:The production of actuate-able materials is an important area of development for Solid Freeform Fabrication (SFF).

Introduction:The production of actuate-able materials is an important area of development for Solid Freeform Fabrication (SFF). Traditionally manufactured actuators have a diverse set of power sources and applications. These range from hydraulic and pneumatic systems, used in heavy equipment, to electric motors, to piezoelectric systems used for small scale manipulations. Each actuator type is generally judged upon many performance metrics, and is then selected for the needs of a particular application. For a 3d printed actuator to be deemed useful, it must be able to compete with its similar non-printed actuators.3D printing of actuators has several potential advantages over traditionally manufactured actuators. 3D printing allows for rapid prototyping, allowing a user to iterate through design very quickly.

3D printing allows for rapid prototyping, allowing a user to iterate through design very quickly. A 3D printable actuator would allow a user to test a design for a new actuator without the costs associated with setting up tooling for a new traditionally manufactured design. 3D printing also offers added geometric complexity over traditional methods. This added geometric complexity may allow for new efficiencies to be achieved. 3D printing can also allow for many diverse material combinations which may allow for complete systems to be produced on a single deviceSolid Freeform Fabrication has had limited success in building actuator systems. Ionic polymer metal composite [IPMC] actuators are generally used in a research context and show promise as artificial muscles and microfluidic valves. They often need to have high reliability, fast response times, and low creep. Dr. Evan Malone and Dr. Jonathan Rossiter have both developed 3d printed IPMC actuators with limited success.

They often need to have high reliability, fast response times, and low creep. Dr. Evan Malone and Dr. Jonathan Rossiter have both developed 3d printed IPMC actuators with limited success. Such actuators tend to be very weak and have short lifecycles, and suffer from many other shortcomings when compared with traditionally manufactured IMPCs. (Malone and Lipson 2005) (Rossiter, Walters and Stoimenov 2009)Attempts to produce complete electromechanical motors using electrostatics and electromagnetic designs have relied on printing the plastic components to a system and manuallyadding wires and conductors. (hakins 2014) (3Ders.org 2014). These are the best efforts today and should not be considered a true 3D printed motor since the vast majority of the complexity and materials in their designs are still produced by hand.

These are the best efforts today and should not be considered a true 3D printed motor since the vast majority of the complexity and materials in their designs are still produced by hand. However they have allowed people to rapidly prototype a design which could be later optimized for manufacture.The most successful electrically powered 3d printed actuators are human muscle cells. While these have no industrial counterpart, they have been integrated into 3D printed constructs to make systems which respond to electrical stimuli. (Cketkovic 2014) Such designs may use added geometric complexity enabled through 3d printing to help the cells receive nutrients in the device while operating.There have been two successful types of 3d printed actuators, piezoelectric and pneumatic. Piezoelectric materials are often used for their high speed movements, small displacements and accurate displacements.

Piezoelectric materials are often used for their high speed movements, small displacements and accurate displacements. These materials have been successfully fabricated using SFF techniques through both direct and indirect methods. (Safari, Allahverdi and Akdogan 2006) 3D printed pneumatics have been produced using SLS systems by Festo and integrated into other complex assemblies. (Festo AG & Co. KG 2013) These actuators have had a high level of reliability and durability. The ability to 3d print pneumatic actuators allows for the integration of many support components (fluid lines, mounting brackets, etc.) to be integrated into a single printed part. Festo’s work and the work on piezoelectric materials has brought the advantages of rapid prototyping to the development of these actuators.Wax actuators are used in a limited set of applications. In general they are used for high reliability short stroke and high power density applications.

In general they are used for high reliability short stroke and high power density applications. Often they are used as part of a temperature regulation system in greenhouses, appliances, HVAC and automotive applications. The system can rely on ambient heat or heat generated by electrical resistance to actuate. A wax actuator traditionally uses a single chamber of wax in a metal enclosure which is attached to a membrane or piston. When the wax melts, it expands causing the piston to move or the membrane to inflate. (Thermomegatech 2014) (Tibbitts 1991) Cooling the wax causes it to contract and return to the original configuration. The melt profile of the wax can be customized to allow for sharp transitions at specific temperatures, or spread out over a range of temperature to allow for positional control between its contracted and expanded state. Solid freeform fabrication has been printing with wax materials for many years.

Solid freeform fabrication has been printing with wax materials for many years. (Solidscape 2014) (Cohen and Lipson 2010) However wax has not been combined with other materials to make a printed wax actuator until this paper.2. Wax actuator Fabrication methods and results: 2.1 EmulationOur early attempts to produce a wax actuator were focused on replicating the design of traditionally manufactured wax actuators. A chamber of ABS was made, and solid wax was placed in the chamber and the top was sealed with silicone. The ABS chamber had been tested for water tightness to assure there were no errors in the print process. When heated the liquid wax would seep out through ABS part and run down the sides. This is most likely due wax having a much lower surface tension than water. This allows the wax to move through very small holes and cracks in the casing.

This is most likely due wax having a much lower surface tension than water. This allows the wax to move through very small holes and cracks in the casing. Attempts were made at coating the ABS parts in silicone to prevent leaking, however this failed to produce a design which would noticeable expand when heated.2.2 Direct Fabrication with dual printingThe second attempt at 3d printing a wax actuator focused on the production of cells of Dow Corning 732 RTV silicone filled with Wax. These were printed on a Fab@Home 3d printer using a pressure based extruder. The top of the cells were left open to produce a structure that was open on the top and sealed on 5 sides. Liquid wax was then filled into the RTV silicone cavity using the robotcasting process, and allowed to sit until the wax cooled and became solid. Silicone was added on top of the cooled cells and allowed to cure.This method resulted in a two problems (See figure 1).

Silicone was added on top of the cooled cells and allowed to cure.This method resulted in a two problems (See figure 1). The wax contracts when cooled causing the top of the system to concave inwards. As a result, the top surface could not be properly sealed since the top surface was not at the proper height. The initial solution was to add additional wax to the cooled wax to fill the voids. The second problem was caused by the low surface tension in the wax. The liquid wax would often coat the top surface of the silicone walls. Since the silicone and wax do not bond, the layers printed on top of these walls would not be able to bond to the wall. Thus when the wax was heated, it would leak from the top of the silicone walls and flow down the sides of the silicone cells.(a)(b)Figure 1: a cell filled with liquid wax (a) develops a depression (b) as it starts to cool.

Overfilling errors and spills can cause wax films to form on the silicone surfaces as seen in the areas highlighted with the blue arrowsThe solution to both problems was to use the print itself as a mold (see Figure 2). The cells would be printed on a Fab@Home Model 3 system and allowed to cure overnight. The mold would then be filled with liquid wax and allowed to cool. Wax would be added after cooling to fill in the depressions left in the center of the cells, and allowed to cool. A blade would then cut off excess wax from the top of the cells. The wax was removed from the cells and set aside. The open and empty cells were then removed from the printer and washed in a bath of water at 85 degrees Celsius for five minutes. This caused the wax which had coated the silicone walls to melt and float to the surface. Manual probing was conducted on the cells to determine that wax had been removed. The cells were then removed from the bath and placed back on the printer.

Manual probing was conducted on the cells to determine that wax had been removed. The cells were then removed from the bath and placed back on the printer. The wax was then placed back into the cells and the top layer of the cells was then printed and allowed to cure overnight.Figure 2: The method for producing wax actuators with Solid Freeform Fabrication and casting contains nine steps. The wax is poured into the printed cell which acts as the mold for the first 5 steps. The cell is cleaned in steps 5 through 8 and finished in step 9.aBFigure 3: a sealed cell after 1 cycle of heating (a) can be reheated, causing expansion in the cell. (b) the cell will find the lowest energy state for the new internal volume. A thin side can direct the expansion into that a direction (b)Closed wax filled cells produced using this method were then measured in the X, and Y directions at their thickest points.

A thin side can direct the expansion into that a direction (b)Closed wax filled cells produced using this method were then measured in the X, and Y directions at their thickest points. They were then put in a bath of water at 85 degrees and left until no solid wax could be felt by manual probing. They were then measured at their thickest point and the change in dimensions were recorded. Cells produced using this method exhibited a 5% to 9% increase dimensional length as a result of heating (see Figure 3).2.3 Novel Material formulationAs an alternative to the method described above, a new material combination was developed. A mixture of a two part Polydimethylsiloxane (PDMS) of Ecoflex 0050 by Smooth- On and wax was used. The mixture produced a suspension of small wax particles inside of a PDMS Matrix. The material could be extruded from a Fab@Home Model 3 System using an 18 gauge tapered plastic tip from Nordson EFD on the Fab@Home system (see figure 4).

The material could be extruded from a Fab@Home Model 3 System using an 18 gauge tapered plastic tip from Nordson EFD on the Fab@Home system (see figure 4). Smaller tips resulted in jams due to the small beads of wax suspended in the PDMS. The size of the beads and distribution of sizes has not yet been determined. The PDMS was selected for its higher elongation at break, lower shore hardness Dow Corning 732 Silicone. Additionally RTV silicone proved unable to easily mix with wax due to its one part acetoxy cure process. The heat from the liquid wax caused the RTV silicone to cure too quickly when mixed and caused the wax to form large chunks, resulting in a material not compatible with robocasting on a Fab@Home systemTo produce the material, the PDMS was mixed together at room temperature, following the procedure described by Smoot-on.

After the ecoflex was mixed, it was placed in an open metal container and liquid paraffin wax was gently added to the metal container so as to prevent material from leaving the container. The PDMS and liquid wax were then mechanically mixed rapidly. The wax cooled as a result of the thermal energy transfer from the wax to the PDMS. The wax was not allowed to pool or stagnate in the container as the entire volume of PDMS and wax was mixed. Mixing was stopped once the entirety of the wax had cooled to room temperature. The mixture of wax and PDMS was then optionally placed in a vacuum chamber to degas the mixture. The PDMS compound has a pot-life of 18 minutes, cures in 3 hours, and must cure at temperatures above 18 degrees Celsius according to Smooth-On. To comply with these requirements, a material batch must be prepared and used rapidly to ensure quality.

To comply with these requirements, a material batch must be prepared and used rapidly to ensure quality. After printing the parts must be allowed to sit to fully cure.In order to characterize the effects of the wax to PDMS concentration, several different ratios of wax to PDMS by liquid volume were produced and tested. In order to test the thermal expansion of the material, molds were out of ABS on an FDM system. The molds were 10 mm wide, 10 mm tall, and 45 mm long. The PDMS-wax mixture was placed in the mold and allowed to cure. The molded PDMS-wax mixture was removed once cured and placed in a bath of water at 85 degrees Celsius, above the melting temperature of the wax. Once all of the wax was melted (as determined through manual probing), it was removed from the liquid and measured in the longest direction using calipers with an accuracy of 0.01mm.Figure 4: The liquid wax (bright orange) is mixed with the PDMS mechanically and rapidly. While mixing the wax solidifies.

While mixing the wax solidifies. The PDMS-wax material must be printed quickly before for PDMS pot-life ends and the viscosity begins to change.(a)(b)Figure 5: The PDMS is clear, but the solid PDMS-wax material takes on a white color.(a) The material becomes translucent to visible light when the wax melts.(b)Expansion in length by wax concentration7.0%h t g n el6.0%5.0%nin o i s n a p x e t n e c r e P4.0%3.0%2.0%1.0%0.0%0%, 0.2%0%10%20%30%40%50%60%70%80%Percent Liquid Wax by volumeFigure 6: Increasing the percentage wax in the material increases the amount the material expands when heated. There is an unexpected decrease in the percentage of expansion between 40% and 50% of wax by volume. The cause of this decrease is unknown.Concentrations of 0%, 20%, 33% 39% 50% and 66% by liquid wax of the material were made and tested.

The cause of this decrease is unknown.Concentrations of 0%, 20%, 33% 39% 50% and 66% by liquid wax of the material were made and tested. As shown in Figure 6 the material showed a definite increase in length relative to the pure PDMS sample. Since PDMS has a relatively high coefficient of thermal expansion of 3.10*10-4 it is expected that the PDMS will increase in size when heated from room temperature to 85 degrees Celsius. Since the ambient temperature was not recorded, it is not possible to compare the results with the value expected in literature accurately. However, an estimate of 21 Celsius for room temperature would give a 0.19% increase in length when heated to 85 degrees Celsius and a 0.2% increase in length was measured. The trend in the data of figure 6, shows that as the concentration of wax is increased, the material expands more. There is a deviation in this trend at the 50% concentration and the cause of this deviation is unknown.𝑚𝑚1𝑚𝑚𝐾3.

There is a deviation in this trend at the 50% concentration and the cause of this deviation is unknown.𝑚𝑚1𝑚𝑚𝐾3. Discussion and future work:Two methods for producing a wax actuator using Solid Freeform Fabrication have been demonstrated. Both the PDMS-wax and silicone cell actuators have several advantages over traditional wax actuators. Both systems are complexly soft, allowing them to be integrated into the growing field of soft robotics. Both system are completely metal free, allowing them to operate in environments like MRIs, where traditional wax actuators would be unable to operate due to their metal housings and pistons. Both systems are volumetrically expanding by nature. This may provide the ideal actuator for systems as described in the evolutionary biology work of Jeff Clune and Nick Cheney. (Cheney, et al. 2013)The direct fabrication method is very time intensive and requires a multistep process for producing a functional wax actuator.

(Cheney, et al. 2013)The direct fabrication method is very time intensive and requires a multistep process for producing a functional wax actuator. The main time constraint on the process, is the curing of the RTV silicone. It has two steps of curing, each requiring 24 hours to ensure a complete bond. Future version of the process should be able to avoid this requirement by using a UV cure silicone or a silicone with a faster cure time. The need for multiple filling steps, a heated bath, and two manual steps make it an inferior process to printing with the novel PDMS-wax material.The PDMS-Wax mixture’s main drawback is the short shelf-life. Using a silicone with a longer pot-life would allow for the usage of the PDMS-wax material to be easier, but would only help by moving the shelf life from minutes to hours. Such longer pot-life materials generally also have a longer curing time, which will only slow down the process of producing wax actuators.

Such longer pot-life materials generally also have a longer curing time, which will only slow down the process of producing wax actuators. Using a UV cured silicone should also allow for an increased shelf life on the order of months or years. Future experiments will have to be done to find a compatible UV silicone material. As seen in figure 5, the PDMS-wax material is milky white when cool and solid but becomes translucent when heated and expanded. This may provide a method of sensing for controlling the actuator in the future.4. ConclusionIn this paper we have demonstrated the first 3D printed wax actuators. We have developed two methods, a direct fabrication method and a novel material for 3d fabrication. The direct fabrication method is a multi-step process which requires manual steps to produce a 3D printable wax actuator.

The direct fabrication method is a multi-step process which requires manual steps to produce a 3D printable wax actuator. The novel material we have developed allows for the direct extrusion of a material which will expand when exposed to ambient temperature above the wax melting point. The methods open up a set of applications for 3d printing in the space of passive thermal management and active thermal actuation.5. AcknowledgementsThe authors would like to acknowledge a grant from the Israeli Ministry of Defense for funding this research.6. Bibliography:3Ders.org. 2014. An 3D printed electrostatic motor with high torque. 7 14.http://www.3ders.org/articles/20131226-free-energy-an-3d-printed-electrostatic-motor-with- high-torque.html.Cheney, N, R MacCurdy, J Clune, and H Lipson. 2013. "Unshackling Evolution: Evolving Soft Robots with Multiple Materials and a Powerful Generative Encoding." Proceedings of the Genetic and Evolutionary Compuation Conference.Cketkovic, C.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed Paper3D PRINTED ELECTRONICSM. Bowa*, M. E. Dean*, and R.D. Horn*Department of Electrical Engineering and Computer Science, The University of Tennessee, Knoxville, TN 37996ABSTRACTAdditive manufacturing is revolutionizing the way we build and produce a plethora of products spanning many industries. It has shown strong potential in reduced energy use, sustainability and cost effectiveness. Exploring avenues that this technology can be utilized is key to improve productivity and efficiencies in various applications including electronic systems and devices manufacturing. Electronic systems and sub-systems are built using a variety of material and processes, which require a large carbon footprint, significant waste material and high production time.

Electronic systems and sub-systems are built using a variety of material and processes, which require a large carbon footprint, significant waste material and high production time. We propose the application of 3D printing technology to support an integrative process for combining circuit board fabrication, solder mask process, electronic component pick and place and enclosure manufacturing. The integration of these separate processes into a single high efficiency additive manufacturing process will yield significant savings in energy use, carbon footprint, waste product and production time and cost.INTRODUCTIONThe electronics industry is growing rapidly with emergence of new technology every day. Additive Manufacturing (AM) has become a game changer in the way we build things and it has immense potential to further integrate into more processes in various industries, specifically the electronic industry.

While the electronics (consumer) industry is a multi-billion dollar industry, it is rapidly growing and so is the demand. Electronic device manufactures are not only looking for cost effective and efficient manufacturing solutions, but ones that will also lessen the carbon footprint, significant waste material, and high production time [1]. Energy efficient solutions and lower production time are not only good for the environment but lower cost, which benefits the manufacturers. So it is imperative to come up with solutions that address most, if not all, of these factors [2]. Advances have and are being made in the drive to find solutions that can be integrated into electronic devices manufacturing using 3D printing, a subset of additive manufacturing. One of these efforts is inkjet printing based. Where conductive nanoparticle liquid inks are used to print traces on 3D printed substrate, plastic or paper and/or encased in 3D printed cases.

One of these efforts is inkjet printing based. Where conductive nanoparticle liquid inks are used to print traces on 3D printed substrate, plastic or paper and/or encased in 3D printed cases. Though this approach meets some of the solutions to the goals stated above, it has challenges. Ink prices and properties such as stability, aggregation and viscosity have to be considered and can be a hindrance to electronic integration that is efficient [3].This manuscript will introduce and serve as a guide that showcases the application of 3D printing technology to support advanced 3D manufacturing of integrated electronic devices using conductive polymer. The current electronic device manufacturing process has many steps that consist of the circuit board fabrication, solder mask process, electronic component pick and place, board soldering/cleaning, and enclosure manufacturing.

This proposed integrative process of2163combining these separate processes into one that will yield 3D printed circuits using conductive polymer filament, and can support digital and analog circuits with an extended level of complexity. When you integrate this capability with 3D automation of component placement and enclosure construction, we believe this research could revolutionizing the way electronic devices like cell phones, TVs, and automotive systems are designed, manufactured, and distributed. All this while saving energy, reducing time from concept to product, reducing carbon footprint and the cost of manufacturing consumer electronic products.MATERIALS AND METHODSTwo 3D printers, the Series Pro 1 and Ultimaker, were selected to perform the necessaryprints for the proposed circuits. The Series Pro 1 has single extrusion while the Ultimaker has dual extrusion, making for more efficient printing.

The Series Pro 1 has single extrusion while the Ultimaker has dual extrusion, making for more efficient printing. Despite the difference in extrusion heads both printers have played an integral role in the printing process.With the printers selected, the conductive filaments that were available to us were evaluated and tested. These filaments consisted of four conductive polymers of which two were Acrylonitrile Butadiene Styrene (ABS) and other two Polylactide (PLA) filaments. The two PLA, a Composite PLA - Electrically Conductive Graphite (ECG) filament from Proto-Pasta and a Conductive Graphene PLA Filament from Black Magic 3D were tested to have the least resistance and were therefore selected for further evaluation. The composite PLA had the following product specification of volume resistivity; 30 Ω-cm 3D printed parts along layers in the X, Y dimension and 115 Ω-cm 3D printed parts against layers in Z dimension (this is a reference to build direction) [4].

And the Conductive Graphene PLA had a product specification with a volume resistivity of 0.6 Ω-cm [5]. Given that resistivity specifications were incomplete due to the fact that the cross-section areas of the materials were not specified, further analysis had to be done.Figure 1 was the model used to further analyze the resistances. The Y represents the 3 different cross-sectional area (width and height), 1.5 x 1.5 mm, 3 x 3 mm, 6 x 6 mm, that were used to determine how the resistance is affected by surface area and volume.

The Y represents the 3 different cross-sectional area (width and height), 1.5 x 1.5 mm, 3 x 3 mm, 6 x 6 mm, that were used to determine how the resistance is affected by surface area and volume. Figure 2 show a 3D printed 3 x 3 mm analysis model.Figure 1: Filament Analysis Model2164Figure 2: 3D Printed Analysis ModelMeasured ResistanceThe first model printed was the 3x3 mm the composite PLA filament, conductivegraphene PLA filament.Table 1: Filament Analysis Results for the Conductive PLAComposite PLA FilamentConductive Graphene PLA FilamentCross-Sectional Area (weight x height)3x3 mm3x3 mmResistance (pad to pad)4.3 kΩ0.13 kΩWith this information, printing and measurements with the other models (1.5x1.5 mm and 6x6 mm) were conducted for both the composite PLA filament and conductive graphene PLA filament. These measurements helped us better understand how resistance varies with the different widths and heights of the models.

These measurements helped us better understand how resistance varies with the different widths and heights of the models. Table 2 below, shows a comparison of the resistances measured for all of the three models of the conductive PLA filaments.Table 2: Filament Analysis Results for the Conductive PLA2165Composite PLA FilamentConductive Graphene PLA FilamentCross-Sectional Area (width x height)Resistance (pad to pad)1.5 x 1.5 mm13 kΩ0.48 kΩ3 x 3 mm4.3 kΩ0.13 kΩ6 x 6 mm1.1 kΩ0.04 kΩAs shown above, as the widths and heights increase, the resistance decreases. Though this new knowledge gave us some insight, we had to further analyze whether resistance varied with volume or surface area.Volume(Vol) and Surface Area(S.A.) in relation to ResistanceBelow are surface area and volume calculations of the analysis model with the three different widths and heights of 1.5 x 1.5 mm , 3 x 3 mm and 6 x 6 mm.

The models length is held constant at approximately 202 mm. Using the 6x6 mm model as reference, a resistance surface area and resistance volume was calculated. Comparing the calculated resistances above with that of table 2, it is observed that the volume resistance factor yielded relatively close resistances to the measured values, compared to that of the surface area resistance factor.

Additionally, observing table 3 below (by the colored text), provides that the volume has a stronger relationship with the resistance and it was therefore concluded that resistance is affected by the volume of the material conducting through the body of the material and not just the surface.Table 3: Filament Analysis Results for the Conductive PLAComposite PLA FilamentConductive Graphene PLA FilamentCross-Sectional Area (width x height (mm)) S.A.(mm2) kΩ.(mm2) Vol (mm3) kΩ.(mm3)1.5x1.5121215756454.559093x324249454181870906x648485333727279991.5x1.51212581.8454.5218.23x32424315.11818236.36x64848193.97272290.92166We can observe the volume resistance calculations in table 3 that the conductive graphenePLA filament shows better correlation between volume and resistance than the composite PLA filament.

This could all be attributed to multiple reasons, which could include; lack of accuracy with the material due to consistency of distribution of conductive particles during and after the printing process, variations caused by print temperature and cooling, size/density of conductive particles in the material, and capacitive load. It’s possible this accuracy can be further improved by increasing the model dimension’s width and heights.The Circuit Printing ProcessWith the eventual goal of implementing both an analog and digital circuit, we focusedon a digital circuit. With reference from a concept, the digital circuit was modeled in Autodesk Fusion 360.

With reference from a concept, the digital circuit was modeled in Autodesk Fusion 360. Figure 3 shows a top view of the 3D circuit design with the electrical components placements specified: a Common Cathode 7 Segment Display, Seven 4.7 kΩ Series Resistors, Four Switches, Four 1.5 kΩ Pull-Up Resistors, 74LS48 BCD to 7 Segment Decoder.Figure 3: The Digital Circuit Layout DesignWith the circuit chosen, it was decided to first concentrate on the digital circuit, given it would be less susceptible to resistance in the traces. Therefore printing analysis of the material to execute suitable pin holes for the components of the digital circuit was done. White nonconductive PLA filament was used as the non-conductive base of all the circuit boards. Figure 4 shows the AutoDesk Fusion 360 3D model for the pin hole analysis.

White nonconductive PLA filament was used as the non-conductive base of all the circuit boards. Figure 4 shows the AutoDesk Fusion 360 3D model for the pin hole analysis. The pin holes had a height of 5mm, outer width of 1.8mm, outer length of 1.9mm, inner width of 0.8 mm and inner length of 0.9mm.2167Figure 4: Pin Holes AnalysisPrinting ChallengesThe composite PLA filament did not offer as many challenges in the printing process asthe conductive graphene PLA filament did. Initially both materials deposited extra material at the start of the print as can be seen in figures 5 and 6. To resolve the extra material deposition at the start of the print (e.g. on the top right corner), a small square pad as well as nonconductive material line traces in between the pin holes were added to the print, as shown in figure 6 and 9. This was done to try and force the extra material to be deposited before the print of the pin holes and to avoid bleeding in between the pin holes.

This was done to try and force the extra material to be deposited before the print of the pin holes and to avoid bleeding in between the pin holes. This however did not solve the problem. Several adjustments to the temperature and columns were tested to achieve better prints. For the composite PLA filament the print temperature was increased slightly above the manufacturers/distributors recommended maximum temperature to 205◦ C. As for the conductive graphene PLA filament, it was found that the suitable printing temperature was 215◦ C. We must also note that the conductive graphene PLA filament material texture caused a slight expansion to the nozzle of the printer from wear, which we had to switch out (detachable nozzles) when we printed with the composite PLA filament.

This affected the print resolution for the conductive graphene PLA filament.Though the printing temperatures helped to a certain degree with the above mentionedchallenges, we still had to take further measures to ensure that the pin holes would print according to the 3D model design. The bottoms of the pin holes were beveled 0.5 1 mm, figure 8 shows the resulting print with the beveling. A line trace was added for extra measure.

The bottoms of the pin holes were beveled 0.5 1 mm, figure 8 shows the resulting print with the beveling. A line trace was added for extra measure. With this result, we were more confident to proceed with printing the digital circuit.2168Figure 5 :Conductive Graphene PLA Filament Pin Holes challenges 1Figure 6: Conductive Graphene PLA Filament Pin Holes challenges 22169Extra material depositionFigure 7: Conductive Graphene PLA Filament Pin Holes adjustments ( with graphene base)Figure 8: Composite PLA Filament Pin Holes challenges 12170Non-conductive material line traces in between the pin holes.Figure 9: Composite PLA Filament Pin Holes challenges 2Figure 10: Composite PLA Filament Pin Holes adjustmentsRESULTS/DISCUSSIONImplementation with both the composite PLA filament and conductive graphene PLAfilament was conducted.

To reduce the resistances across the circuit, with the information learned in filament analysis, without compromising the circuit design, the line traces were designed with a height on 4mm and width of 0.5 mm, while the pin holes remained the same dimensions shown in figure 4. A height of 5mm, outer width of 1.8mm, outer length of 1.9mm,2171inner width of 0.8 mm, inner length of 0.9mm, and gaps between the the pin holes was 0.7mm. With the gap between the pin holes being 0.7 mm, we were unable to place any traces in between the pin holes due to the print resolution.Figure 11: Composite PLA Digital Circuit PrintWe experienced no issues with printing the digital circuit with the composite PLA filament.As can be seen in both figures 11.

The design was adjusted to extend the ground line trace on the right side to create another ground connect in the hopes of further reducing resistance in the ground circuit.Figure 12: Conductive Graphene PLA Digital Circuit PrintThe conductive graphene PLA filament line traces were also designed with a height on 4mm, width of 0.5 mm, and the pin holes remained the same dimensions shown in figure 4. The digital circuit print for this filament, as a whole, was relatively good. However two spots in the circuit line trace was ragged in certain layers of the print. This could have been due to2172clogging and the material did not deposit evenly. We mitigated the issue by thickening the traces 0.5mm or less, which seemed to alleviate the problem.Circuit TestsAfter the circuits were printed, resistance of the circuit traces were measured to observeresistance across the circuit at various points.

Using figure 13 as reference and the resistance measurement lists below, R(1-2) represents the measurement between points (1) and (2). R(23) represents the measurement between points (2) and (3). R(I.(4)), R(II.(4)), R(III.(4)), R(IV.(4)), R(V.(4)), R(VI.(4)), R(VII.(4)) represents the measurements between the points identified (4) in figure 13. R(5-5), R(6-6), and R(7-7) represents the measurement between point (5) and (5), (6) and (6), and (7) and (7), respectively. As displayed by the resistance measurements listed below (from the different points of measurement) that the conductive graphene PLA filament circuit measured lower resistance than the composite PLA filament circuit.

This is displayed in figures 14 and 15. It can also be seen that there is an obvious response from the circuit, in terms of connectivity, however the display did not switch numbers when prompted by the switches.2174Figure 14: Composite PLA Digital Circuit PrintFigure 15: Conductive Graphene PLA Digital Circuit PrintThere various possible reasons as to why the both circuits, shown in figure 14 and 15,did not produce any other results when promoted by the switches. It could have been due to different resistance the traces showed as seen in figure 13, causing poor conductance in certain areas of the circuit. Furthermore, the ragged traces seen in the conductive graphene PLA filament print, definitely created additional resistance.

Furthermore, the ragged traces seen in the conductive graphene PLA filament print, definitely created additional resistance. Finally, we have not yet looked into a relationship between the material and its capacitance and/or inductance, but it could have also played a role in impeding the flow of current in some of the line traces.2175CONCLUSIONFrom the analysis and initial testing, a 3D printed digital circuit was successfullyimplemented using the composite PLA filament and conductive graphene PLA filament, as shown in figures 14 and 15.

There are still a few adjustments that need to be made to attain optimal functionality, however, this analysis and initial testing highlights tremendous promise that a digital circuit can in fact be 3D printed with the both filament materials.Moving forward, a better characterization of the materials needs to be done for consistency.A further look into the material make up, as in what amounts of graphene and graphite are present in the materials and how it relates to the conductivity. This information was not provided by the manufacturers/distributors and further research into the materials must be conducted. In addition to future work, there are plans for implementation and testing of analog circuit with these conductive polymers.2176REFERENCES[1] industry in 2017. [Online]. Available: https://www.pannam.com/blog/ top-trends- andchallenges-in-electronics-manufacturing/Hooper, T. (2017).

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference3D Printed Fastener-Free Connections for Non-Structural and Structural Applications – An Exploratory InvestigationDaniel Delgado Camacho*, Patricia Clayton*, William J. O’Brien*, Kee Young Jung*Department of Civil, Architectural and Environmental Engineering, The University of Texas at Austin, USAAbstractThe construction industry has shown increasing interest in AM technologies and has successfully implemented various proof of concept projects using different AM processes. Much of the research on AM in the construction industry has focused on development of new large-scale extrusion printing systems and on development of cementitious materials for AM applications, whereas research exploring new applications of already existing AM technologies and materials suitable for construction applications has been scarce.

This paper explores the use of existing, small-scale material extrusion 3D printers to create fastener-free connections that could be used in structural or non-structural applications. These connections, inspired by traditional wood joinery and modern proprietary connections were printed using polylactic acid (PLA) material. The flexural strength of the connections was then tested using a four-point bending test to evaluate their potential structural performance and to identify connection types that warrant further research in this exploratory proof of concept study.Keywords – Fastener-Free Connections, Additive Manufacturing, 3D Printing, Material Extrusion, Polymers, Flexural Test1. IntroductionAdditive manufacturing (AM), the process of building up components in a layer-wise fashion from 3D models, has been used increasingly in many industries such as automotive and aerospace.

AM seven different manufacturing processes: vat photopolymerization, material jetting, binder jetting, material extrusion, powder bed fusion, sheet lamination, and direct energy deposition. These processes are explained in more detail in the ISO/ASTM 52900-15 Standard [1]. Current AM technology is generally used in smaller scale applications compared to the scales encountered in infrastructure construction projects. When wanting to apply AM technologies to scales that are more common in construction, two major limitations, print volume and print time, become a challenge. The construction industry has shown interest in the use of more automated AM technologies and has started exploring proof of concept applications using different AM processes, mostly material extrusion processes using cementitious materials [2].

AM could be utilized in many ways in the construction sector such as producing novel forms, optimized structural topologies, customized parts, in situ repair, and tolerance matching [2]. Interdisciplinary research is still needed to make this technology reliable at larger scales and for construction applications.is categorizedintoIn the construction industry, the most common types of materials used are wood, steel, and concrete. Wood is commonly used for low-rise residential construction and for formwork.580Subtractive processes are often used to cut wood to specified dimensions and details either manually or using CNC machines, depending on the application. Steel is commonly used as a main structural component in industrial and high-rise applications, where structural components are typically fabricated using a combination of formative and subtractive processes (e.g., casting, rolling, cutting, drilling).

Fabrication of reinforced concrete structural components requires the use of formwork, typically made from wood or steel, to shape the wet cementitious material after mixing and casting until it finishes curing. Both subtractive and formative manufacturing are common methods used in the construction industry. As it has done in other industries, additive manufacturing can provide an alternative construction process when challenges are present using these other two methods. For example, manufacturing reinforced concrete elements that have unique dimensions would require changing the mold/formwork or the milling tools used, which can result in prolonged manufacturing time, increased manufacturing cost, and potentially delayed delivery/construction. The use of AM could allow the designer and erectors to customize a design based on site-specific constraints without having to rely only on long-standing manufacturing and construction processes.

The use of AM could allow the designer and erectors to customize a design based on site-specific constraints without having to rely only on long-standing manufacturing and construction processes. For these reasons, AM has become an attractive option when dealing with unique single production needs. AM also has been used as a prototyping tool to check final design and assembly before mass production or before production at larger scales.Much of the research on AM in the construction industry has focused on development of new large-scale extrusion printing systems and on development of cementitious materials for AM applications, whereas research exploring new applications of already existing AM technologies and materials suitable for construction applications has been scarce. As advancements continue to be made in AM processes and materials, there are likely to be more structural and nonstructural AM components used in construction.

As advancements continue to be made in AM processes and materials, there are likely to be more structural and nonstructural AM components used in construction. These components may include architectural claddings [3], formwork for concrete components of complex geometry, permanent or temporary structural elements such as beams and scaffolding, and much more. Even as large-scale AM technologies allow for full-scale components, and even monocoque structures, to be printed, it likely that multiple structural components and architectural finishing will be printed on- or off-site and will need to be connected to form the final product.

While traditional connection methods using fasteners such as bolts, screws, welds, epoxies, or un-bonded post-tensioned elements [4], can be used to connect these AM components, there may be novel connection methods amenable to AM that have otherwise considered impractical or uneconomical in conventional construction practices.This paper explores the use of AM to create fastener-free structural or non-structural connections of complex geometry that are well suited for AM processes over conventional formative and subtractive processes used in construction. This exploratory research focuses on proving and testing the connection concepts at small scales and with commercially available 3D printers, such that promising connections can be identified and later explored at larger scales and with other materials. The connections in this study are inspired by traditional woodworking joinery and are manufactured using a material extrusion process.

The connections in this study are inspired by traditional woodworking joinery and are manufactured using a material extrusion process. The flexural strength of the connections is evaluated using a four-point bending test. Potential uses for these fastener-free connections in construction could be for assembling formwork, for attachment of architectural components, rapid erection scaffolding, to join beams to columns, or as beam splices. The results of this study can inform future research that investigates the most promising connections using different scales, materials, and other AM processes and technologies.5812. Inspiration for Fastener-free ConnectionsErection and assembly of formwork and prefabricated structural and nonstructural components are some of the most time-consuming tasks on a construction site [5]. The time needed to make on-site corrections due to misfits and tolerance issues can further complicate and delay construction projects.

The time needed to make on-site corrections due to misfits and tolerance issues can further complicate and delay construction projects. Therefore, recent research in construction has focused on facilitating more rapid construction processes and more flexible on-site tolerance accommodation. These methods of construction have typically focused on improving efficiency using familiar construction techniques. For example, to reduce on-site construction time, modular structural or nonstructural subassemblies can be prefabricated off-site; reusable steel formwork is used to produce precast concrete off-site or for repeated cast-in-place concrete components. For each of these methods, pre-fabricated components or subassemblies must be assembled and connected on-site, typically using fasteners such as nuts and bolts, screws, nails, adhesives, and more.

For each of these methods, pre-fabricated components or subassemblies must be assembled and connected on-site, typically using fasteners such as nuts and bolts, screws, nails, adhesives, and more. This research aims to explore “quick connection” concepts that do not rely on conventional fastener techniques but instead rely on geometries that can be realized using additive manufacturing processes.Inspiration for these “quick connection” concepts was drawn from Japanese wood joinery and connection concepts from modern proprietary companies such as ConXtech Inc. and K’NEX. ConXtech Inc. connections are intended for structural applications in modern steel construction, while K’NEX serves more as an educational toy concept to enhance children’s creativity. These types of connections rely on steel CNC milling and plastic injection molding, respectively, to produce connection geometries capable of transferring loads.

These types of connections rely on steel CNC milling and plastic injection molding, respectively, to produce connection geometries capable of transferring loads. Japanese wood joinery, on the other hand, relies on historic woodworking techniques that have been passed down from generation to generation. In Japan, skilled craftspeople dedicated to connecting elements of a structure were considered master joiners [6]. The joints had to be able to transfer forces between members, typically without the use of fasteners, and at the same time had to be aesthetically pleasant [6]. This type of joinery is seen as pieces of art, and unfortunately difficult and time-consuming to reproduce by hand. Subtractive processes such as using CNC machines can help to facilitate fabrication of these complex joints, but some of the fine details of the connection are lost due to CNC tooling limitations.

Subtractive processes such as using CNC machines can help to facilitate fabrication of these complex joints, but some of the fine details of the connection are lost due to CNC tooling limitations. Using these types of fastener-free connections as inspiration, this research investigated types of fastener-free connections for structural or nonstructural applications that rely on complex geometries suitable for AM and that would otherwise be difficult to create using conventional subtractive and formative processes commonly found in construction. The three connection concepts that are explored in this study are the Kawai-tsugite, Kanawai-tsugi, and ConXtech connections.2.1 Kawai-tsugiteThe Kawai-tsugite connection [7, 8, 9, 10, 11] was created in the 1980s by Professor Naohito Kawai when he was a student at the University of Tokyo [10].

Kawai-tsugite is not a typical joint in Japan due to its complex geometry [11], and it was created primarily to test woodworkers’ craftsmanship skills; thus, information and details about this specific joint are lacking in Japanese joinery books [11]. In the name, Kawai comes from the creator of the joint, Naohito Kawai, and tsugite is the name commonly given to describe a splice joint in Japanese.Kawai-tsugite uses cube symmetry to join itself in three different ways, shown in Figure 1. A cube has a rotational symmetry of 120° about the axis going through opposite corners of the582cube, meaning that when a cube is rotated 120° about this diagonal axis, the cube will appear to maintain is original geometry. Kawai-tsugite uses geometry in a complex way that is difficult to create or explain; thus, more details on fabrication of this connection can be found at [12].Figure 1.

Kawai-tsugite uses geometry in a complex way that is difficult to create or explain; thus, more details on fabrication of this connection can be found at [12].Figure 1. Assembly options for Kawai-tsugiteFigure 2 shows the dimensions used for the main connection based on a 37 mm x 37 mm x 37 mm [1.5 in. x 1.5 in. x 1.5 in.] cube. The two pieces being connected are geometrically identical but can be rotated in different orientations to produce splice joints and right-angle joints in different directions. This connection concept was selected for this study due to its complex geometry that is well suited for AM but is difficult to produce using conventional subtractive processes.25mm12mmPROJECTTITLEDRAWNCHECKEDAPPROVEDSCALEWEIGHTSHEETDWG NOREVCODESIZE1/11:205/04/2018Daniel Delgado CamachoA Additive Manufacturing UTKawai Tsugite for Drawing12mm25mmFigure 2.

Kawai-tsugite concept joint2.2 Kanawa-tsugiKanawa-tsugi is a splicing joint, described as a mortised rabbeted oblique splice by Suimyishi and Matsui [6]. There are many joints in traditional Japanese joinery that are similar to this one, such as the sashimono, shiribasami-tsugi, or the shippasami, with the main differences between each being small changes in the dimensions, angles, and layout of each part of the connection. This connection shows two identical pieces that fit together and slide parallel to each other to lock into place. This motion and dimensions of the connection leave a small gap (shown583in Figure 3) in the middle, between the two pieces after they are locked into place. A pin is inserted in this small gap using friction to secure the pin and the connection.

A pin is inserted in this small gap using friction to secure the pin and the connection. This type of splice, when made from wood, has an ultimate tensile strength higher than other traditional Japanese splice joints [6], hence its selection to be included in the current study for applications in construction. Figure 4 shows the Kanawa-tsugi connection used in this study.Figure 3. Kanawa-tsugi assembly sequence13mm13mm38mm22mm13mm6mm57mmPROJECTTITLEDRAWNCHECKEDAPPROVEDSCALEWEIGHTSHEETDWG NOREVCODESIZE1/11:205/04/2018Daniel Delgado CamachoA Additive Manufacturing UTShippasami Tsugi Final 0.005 Tol USB 2 (not blue color) v16mm19mm13mm22mm6mmFigure 4. Kanawa-tsugi joint concept2.3 ConXtech and K’nexThe final joint explored is inspired by connections produced by the company ConXtech Inc., which focuses on rapid and efficient modular construction using customizable steel connections for beams and columns [13].

These proprietary connections utilize tapered collar584geometries to create gravity stabilized connections that can be “lowered and lockedTM” in-place and can be used to assemble beams and columns faster and safer, shown in Figure 5 [13].Figure 5. ConXR200 connection [11]Bolts can be added to these connections as needed for structural resilience. Even though three different connection styles called ConXRTM, ConXLTM, and the company has ConXGravityTM, the connection in this study was primarily inspired by the steel moment-resisting frame ConXRTM connection. These connections are manufactured using subtractive processes such as CNC milling and using robotic welding, with predetermined dimensions that can be selected by the structural engineer [13]. A similar connection concept with a very different application came from K’NEX [14].

A similar connection concept with a very different application came from K’NEX [14]. This company focuses on creating educational toys for kids with a kit of parts that can be used to create buildings, bridges, roller coasters, and more. The kit of parts comes with multiple connectors and stick-type elements that could be disassembled and re-assembled as desired. The K’NEX concept of reusable and standardized kit of parts inspires future sustainable construction and design practices where buildings can be built, disassembled, reconfigured, and reused for future applications as needed. This connection concept was selected for this study due to its familiarity and use in modern construction and educational applications.Figure 6 shows a 3-D model of this ConXtech/K’NEX inspired connection.

The connection resists flexural demands through the mortise and the tenon, where the mortise is considered the reduced section or cavity that will receive the tenon, and the tenon is the portion of the joint that protrudes from the circular stub in the middle of the joint. A solid side was added to the tenon element to prevent rotation or twisting of the connection once in place.38mm14mmMortiseTenon14mmPROJECTTITLEDRAWNCHECKEDAPPROVEDSCALEWEIGHTSHEETDWG NOREVCODESIZE1/11:205/04/2018Daniel Delgado CamachoA Additive Manufacturing UTConXtech Final 0.005 TolFigure 6. ConXtech concept joint5853. Background on 3D Printing3.1 3D Printing ProcessThe 3D models of the connections shown in previous figures were created using Autodesk® Fusion 360TM software [15].

ConXtech concept joint5853. Background on 3D Printing3.1 3D Printing ProcessThe 3D models of the connections shown in previous figures were created using Autodesk® Fusion 360TM software [15]. Models for AM applications can be created using a wide range of commercially available 3D modeling software such as SketchUp software [16], SolidWorks software [17], Inventor® 3D CAD software [18], AutoCAD® 3D software [19], 3ds Max® software [20], etc. Users interested in 3D printing can choose the software based on their expertise and preference. Autodesk® Fusion 360TM is a cloud-based 3D computer aided design, manufacturing, and engineering (CAD/CAM/CAE) software [15]. Fusion 360 allows users to convert a 3D model directly into a .stl file, which is the standard format for 3D CAD models that can be used in 3D printing software to slice the part into printable layers [21].

While many options are available, slicer software is typically developed to work best with specific printers and specific applications. Once the .stl file is imported to the slicer software, the user orients and sizes the model as needed. Inside the slicer software the user needs to specify the material being used. The most common materials for 3D printing or rapid prototyping are Acrylonitrile Butadiene Styrene (ABS) and Polylactic Acid (PLA). Other settings that need to be specified are: the extrusion temperature of the material, extrusion rate, nozzle travel speed, the layer thickness, infill ratio of the model, if the print will require supports, and many other case specific settings. Once those settings are specified, the software slices the model based on the layer thickness inputted and creates a .gcode file that contains all the information needed for the 3D printer.

Once those settings are specified, the software slices the model based on the layer thickness inputted and creates a .gcode file that contains all the information needed for the 3D printer. A .gcode file will send a set of commands one by one to let the printer know where and how much material will be extruded and the specific travel pattern, temperatures, and speed of the nozzle. G-codes are already used in subtractive processes such as CNC to automate machine tools and specify where to remove material from a solid block. The same concept is used for AM, but instead of removing material, AM adds material in a layer by layer fashion. More details about AM processes can be found in Gibson, Rose, & Stucker [21].3.2 3D printing Design Considerations3D printing is commonly used for rapid prototyping to test ideas, dimensions, and assembling processes before submitting final drawings or models for mass production.

Rapid prototyping allows designers to present their concepts to others in a more realistic 3D format rather than simply showing 3D computer models or 2D drawings. Having a simple 3D prototype allows the audience to better evaluate and provide feedback on a design for future iterations. Prototypes can also be used for testing to quantify relative performance between different part geometries, which is the prototyping application being used in this study.For construction, tolerances are an important factor in fabrication and performance, where even a fraction of an inch of difference in a part’s geometry can significantly increase construction time required to make modifications necessary for the multiple parts to fit together.

For example, the layout of structural concrete and light-frame wood construction components are suggested to be within ±6.4 mm [1/4 inch] of specified locations [22, 23], and any finishes or additional components connecting to these elements must be fabricated to accommodate these tolerances. For subtractive manufacturing processes, the tools used to carve material out of a solid block will govern the possible tolerances. For formative processes, tolerances depend on the formwork586tolerances and correct placement. Due to the automation of 3D printers, parts can be manufactured to relatively tight tolerances, from micron to millimeter tolerances depending on the print scale [21], compared to the tolerances of about ±6.4 mm [1/4 inch] for hot-rolled steel cross-sections traditionally used in construction per the American Institute of Steel Construction [24].

In this study, gaps were provided between the connecting parts to accommodate fabrication tolerances, and different gap sizes were explored in initial connection prototypes. Smaller gaps made connection fit-up difficult, while providing larger gaps made the connection lose the capability to stay in place. Ultimately, a gap of 0.508 mm [0.02 in] was provided between the parts in the connection. Comprehensive exploration of gap tolerances for conventional or 3D printed connections could be a research topic in and of itself and is thus considered outside the scope of this study. Potential research could study the effects of gap tolerance size as the scale of the printed part is increased or decreased. For example, if larger layer thicknesses are used in AM, by how much, if at all, should the connection tolerances be increased?For AM, different factors might affect the quality of tolerances.

For example, if larger layer thicknesses are used in AM, by how much, if at all, should the connection tolerances be increased?For AM, different factors might affect the quality of tolerances. Since AM is a layer wise approach, the layer thickness will dominate the quality of the final print, especially for finer details. In the material extrusion process, a tabletop 3D printer typically uses a layer thickness of 0.2 mm [0.008 in.] with a standard nozzle size of 0.4 mm [0.016 in] diameter for a “normal” quality print [25, 26]. Higher quality will require smaller layer thicknesses that might also be able to meet stricter part geometry tolerances. Material extrusion processes are generally only able to provide reliable prints with a layer thickness of 0.1 mm [0.004 in.] or larger [26], compared to other AM processes such as vat photopolymerization that can give a layer height as low as 0.025 mm [0.001 in.] [27].

A drawback from producing prints with smaller layer thicknesses is the increase of the number of layers required, which increases printing time. The increased number of layers might also increase the number of weak points or slip planes in between the layers that could induce premature failure. As an example, smaller layer thicknesses have been shown to induce larger thermal strains that cause thermal stresses in 3D printed components [28]. 3D printing is evolving rapidly and potential improvements in print processes are readily available. With the advantage of being open source technology, users can make modifications to their machines to improve their 3D printing experience and can share their knowledge to the 3D printing community.

With the advantage of being open source technology, users can make modifications to their machines to improve their 3D printing experience and can share their knowledge to the 3D printing community. An example of this community-based improvement is the recent ability in some slicer software to change the layer thickness within an individual part depending on the quality that you want to get at different locations [29].As mentioned before, the volume of 3D printed parts allowed in commercially available tabletop printers is relatively small, around 300 mm x 200 mm x 440 mm [12 in. x 8 in. x 17 in.] [25] for the tabeltop printers used in this study, compared to the volumes you expect on a construction site. Although there is research focused on large scale printing, such as the work being done at Oak Ridge National Laboratory (ORNL) [30], small scale additive manufacturing is the most ubiquitous 3D printing technology currently available.

Thus, this proof-of-concept paper on 3D printed connections focuses on smaller size connections printed using commercially available tabletop printers. As 3D printing technology advances and limitations in print scale and print time are addressed, fastener free connections identified in this study to have promise for structural and nonstructural applications can later be investigated at larger scales.5874. Evaluation of 3D Printed Connections4.1 3D Printer and SettingsFor this investigation, the Longhorn® Maker Studio facilities at the University of Texas at Austin were used for prototyping and final production of the different connections that were tested [27]. A CraftBot XL, which has a build volume of 300 mm x 200 mm x 440 mm [12 in. x 8 in. x 17 in.] [25], was used for this study. The material chosen was Hatchbox orange PLA 3D printer filament, with a dimensional accuracy of +/- 0.03 mm [0.0012 in.], and a diameter of 1.75 mm [0.069 in.].

The material chosen was Hatchbox orange PLA 3D printer filament, with a dimensional accuracy of +/- 0.03 mm [0.0012 in.], and a diameter of 1.75 mm [0.069 in.]. PLA is a biodegradable thermoplastic commonly used in the 3D printing community due to its ability to heat and print with accuracy [32]. Another material that is commonly used for 3D printing is ABS, which is a petroleum-based thermoplastic with higher flexural strength, ductility, and durability [33, 34]. The challenge with ABS is that is more sensitive to temperature changes and produces fumes [34]. PLA was chosen as a starting point for this research as it is readily available and a commonly used material in tabletop 3D printers. Future research could explore the variability of mechanical properties and failure modes when different materials are used, as well as the limitations of using the other materials.The settings used for this printer were the default “easy mode” settings by the Craftware software.

The platform bed and the printer head were heated to 60 °F and 215 °F, respectively. The build plate was also coated with hair spray to improve adhesion of the initial printed layers. A layer thickness of 0.2 mm [0.008 in.], considered as a high print quality for fused filament fabrication (FFF), was chosen. The travel speed was reduced from 180 mm/s [7.9 in./s] to 120 mm/s [4.72 in./s]. The infill speed was reduced from 150% to 100%. The reason to reduce the travel speed and infill speed was mainly to avoid issues like layer offsets during printing, which can often occur due to high speeds during printing. The type of infill selected was a square grid, which was the default, and the density was set to 70% infill. The infill pattern and density was kept constant for all prints in this study, and was therefore not investigated as an experimental variable. A 70% infill was used instead of a solid print (100% infill) to decrease material use and printing time.

A 70% infill was used instead of a solid print (100% infill) to decrease material use and printing time. Using CraftBot XL to print specimens with the dimensions shown in Figure 7 with the specified settings took around 19 hrs. to print each specimen.207mmPROJECTTITLEDRAWNCHECKEDAPPROVEDSCALEWEIGHTSHEETDWG NOREVCODESIZE1/11:205/04/2018Daniel Delgado CamachoA Additive Manufacturing UTSolid Beam31mm31mmFigure 7. Dimensions for specimens5884.2 4 Point-Bending Flexural TestA four-point bending flexural test was used to explore the flexural capacity and the failure mode of each type of fastener-free connection. As illustrated in Figure 8a, four-point bending test consists of placing a beam on bottom supports spaced at a length of L and with top loading points spaced at L/2. A constant load rate is applied to the specimen until it fails, where flexural failure is expected to occur in the middle half of the beam where the bending moment is largest.

A constant load rate is applied to the specimen until it fails, where flexural failure is expected to occur in the middle half of the beam where the bending moment is largest. As shown in Figure 8b and Figure 8c, the ends of the beams, between the supports and loading points, is where shear is non-zero and the middle portion of the beam exhibits a constant bending moment at its maximum value. Thus, the connections under investigation were placed in this middle part of the beam specimen, to test their capacities in pure bending. ASTM D6272 -17 [35] provides guidance to determine the flexural properties of unreinforced plastics via four-point bending tests, assuming the tested part is made of completely solid and isotropic material. The FFF material extrusion process is a layer by layer process, resulting in anisotropic material properties; thus, results are expected to be different than those assumed in the ASTM standard.

The FFF material extrusion process is a layer by layer process, resulting in anisotropic material properties; thus, results are expected to be different than those assumed in the ASTM standard. Additionally, the additively manufactured specimens being tested have a specified 70% infill, which is not consistent with the solid infill assumed in the current standards.Figure 8. Shear and moment diagram for four-point bending flexural test5895. Test ResultsThe results of the four-point bending tests for the reference beam specimens (with 70% infill and no connection) and specimens with the three different types of connections (Kawai- tsugite, Kanawa-tsugi, and ConXtech) are summarized below. At least three specimens were tested of each type, as indicated in the specimen naming convention by the number following the specimen type.

At least three specimens were tested of each type, as indicated in the specimen naming convention by the number following the specimen type. All specimens were printed using the CraftBot XL printer with the same environment conditions in the Longhorn® Maker Studio facilities at the University of Texas at Austin. Thermoplastics are greatly affected by temperature difference in the room in which printing is done, which could cause some shrinkage of the material and cause warping of the 3D printed part. It is worth noting that all specimens for this project were printed from December 2017 to January 2018, when University classes were not in session and building thermostats were set to lower temperatures to conserve energy; thus, the researchers were unable to control the ambient room temperature. Some minor shrinkage was observed in the specimens.

Some minor shrinkage was observed in the specimens. Other printing issues that arose with individual connection types will be discussed below, where applicable.5.1 Reference BeamTo serve as a baseline for the capacity of the fastener-free connections, reference beams, with 70% infill and no connections at midspan, were printed and tested. The testing machine used required an initial pre-load recommended to be no less than 445 N [100 lbs.] by the technicians. This preload is to ensure the loading fixture is engaged with the specimen. Figure 9 shows the load vs displacement of the reference beams. The maximum loads on these beams were around 9200 N [2090 lbs.] with a max displacement of roughly 3.1 mm [0.12 in.] and they exhibited brittle failure. Figure 10 shows the failure mechanism of a reference beam after it reached its peak capacity.

Figure 10 shows the failure mechanism of a reference beam after it reached its peak capacity. The failure at the midpoint of the beam began at the bottom of the cross-section where the tensile stresses were largest and propagated up towards the top of the section.Figure 9. Load vs. displacement response for reference beams590Figure 10. Reference beam after testing5.2 Kawai-tsugite ResultsSpecific issues that were observed while printing this connection was stringing (shown in Figure 11) when the nozzle changed path causing strings of polymer to be left in places where it was not intended. Mill files were used to remove such imperfections in the specimens, which could affect the tolerances and fit-up of the printed parts.

Mill files were used to remove such imperfections in the specimens, which could affect the tolerances and fit-up of the printed parts. Stringing typically occurred in the side of the printer that was closer to the door and is believed to be caused by the colder temperatures on this side of the printer due to the very low room temperature outside of the printer (note, the specimens were printed when the university was on holiday break, and thermostats were set lower to conserve energy).The excess material from the stringing was removed and tested, and no changes were made to the printing process to mitigate stringing.Figure 11. Stringing occurring during printingFigure 12 shows the load vs. displacement relationship for the Kawai-tsugite connection. In Figure 12, an initial constant load applied to the connection can be observed, resulting in significant displacement of about 1 mm to 1.5 mm [0.04 in. to 0.06 in.] before the specimen begins to resist load with higher stiffness.

to 0.06 in.] before the specimen begins to resist load with higher stiffness. It was concluded that such displacement might be due to the gap provided in the connections to accommodate tolerances, where an initial displacement is required for the connection to engage and resist the applied load. Kawai-tsugite 2 began to resist load at a smaller displacement, which is believed to be due to the excess material caused by stringing that occurred during printing, which, even after being removed, might have reduced the gap in the connection, causing the connection to engage earlier. The maximum load that this connection was able to resist was around 2400 N [540 lbs.] with a maximum displacement of roughly 4.4 mm [0.17 in.], which is consistent with the other Kawai-tsugite specimens.591Figure 12. Load vs.

Load vs. displacement response for Kawai-tsugite connectionFigure 13 shows pictures of the connection before and after testing, where the layer thickness is 0.2 mm [0.01in.] and the small gap of 0.508 mm [0.02 in.] between the printed parts can be seen to accommodate printing tolerances. Figure 13b and Figure 14 reveals the failure mechanism of this connection. A crack initially formed parallel to the layer orientation (perpendicular to the load) and continued growing perpendicular to the layer orientation until it failed in a sudden, brittle manner.(a)(b)Figure 13. (a) Before and (b) after testing592Figure 14. Four-point bending test for Kawai-tsugite connection from two different angles5.3 ConXtech ResultsA total of four ConXtech specimens were printed and tested; however, the ConXtech 3 specimen encountered several problems during printing, which ultimately affected its behavior and prompted the researchers to print and test a fourth specimen of this type.

During printing, the ConXtech 3 specimen exhibited warping (an example of warping is shown later in Figure 19a), which is believed to have reduced the tolerance gap in the connection, causing the connection to engage earlier in the loading protocol. The warping is believed to be caused by the temperature differential between the printing environment and the surrounding room temperature, as previously described. Due to the cold room temperature, the printed parts, particularly ConXtech 3, did not properly stick to the print bed, causing subsequent layers to be detached and warped. In addition to the warping, other problems occurred during printing that are commonly observed in the 3D printing community. An example is a layer shift as shown in Figure 15, which is caused by the nozzle traveling too quickly, causing it to skip some steps and produce subsequent layers that are no longer aligned with the specified coordinates for the previous layers.

ConXtech 1 and 2 were printed on the same printer as the Kawai-tsugite specimens, but ConXtech 3 and 4 were printed on a different printer of the same CraftBoxXL make and model that is referred to herein as Printer 2. Ultimately, ConXtech 4, which was printed without issues, exhibited behavior similar to ConXtech 1 and 2, indicating that the change of printer did not have a significant effect on performance when issues were not encountered during printing.Figure 15. Layer shifting593Figure 16 shows the load vs. displacement relationship for the ConXtech connections. The maximum load that the connection could sustain was around 2,500 N [560 lbs.], similar to Kawai- tsugite. The connection starts resisting load after displacing approximately 1.3 mm [0.05 in.], when the connection engages.

The connection starts resisting load after displacing approximately 1.3 mm [0.05 in.], when the connection engages. A major difference of this connection is that it had a more ductile failure compared to Kawai-tsugite, as indicated by the reduced stiffness and increased plastic deformations prior to failure, or loss of load carrying capacity.Figure 16. Load vs. displacement response for ConXtech connectionFigure 17 shows before and after testing photos of a ConXtech connection, where you can see the two different parts of the connection. The piece on the left in Figure 17a can be considered the mortise, since it has a void to receive a tenon. The tenon has a projection that is inserted to the mortise. In Figure 17 and Figure 18, the failure mechanism can be observed where the tenon fails in tension where its cross-sectional area is smaller.

In Figure 17 and Figure 18, the failure mechanism can be observed where the tenon fails in tension where its cross-sectional area is smaller. It can also be seen in Figure 17b and Figure 18 that the mortise formed a crack at the middle of the connection where it started to fail parallel to the layer orientation.Figure 17. (a) Before and (b) after testing594Figure 18. Four-point bending test for ConXtech from two different angles5.4 Kanawa-tsugi ResultsSeveral issues that are common to 3D printing were observed while printing the specimens for the Kanawa-tsugi connection, such as, bed leveling issues, electrical connection issues that caused the printer to stop, temperature differences causing the printed part to warp or not stick to the bed, and the nozzle hitting the printed part and causing it to fail.

Due to the several problems that were encountered, it was suggested by [36] to add a raft (shown in Figure 19b) at the bottom of the specimen to increase attachment to the bed and decrease warping. The raft is removed prior to testing. Another suggestion was to print each piece of the connection separately to avoid failure of the entire print should an issue arise. It is believed that these differences in the printing process used for these specimens (i.e., the addition of the raft during printing and printing the different connection parts separately) did not significantly affect connection behavior, since the added material was removed before testing. Kanawa-tsugi specimens had the greatest variability of results and the lowest capacity of all connection types, which is due to the reduction of effective cross-sectional area resisting the flexural demands within the connection.

It is unclear whether the variability in response of the Kanawa-tsugi specimens is due to the connection itself or due to variability in the printed part. The depth of the effective cross-sectional area in the connection was a third of the gross cross-sectional area, reducing the moment of inertia compared to the other two connections, thus, reducing its moment capacity and maximum load.(a)(b)Figure 19. (a) Example of warping, and (b) an added raft to decrease warping595Figure 20 shows the load vs. displacement relationship for the Kanawa-tsugi connection. Consistent with the responses of all the other specimens, a constant load is observed at the beginning of the test where the specimen displaces before the connection engages and begins to resist more load. The maximum load for this was about 1,600 N [360 lbs.] with a maximum displacement of 3 mm [0.12 in.]. All the Kanawa-tsugi connections were printed on a 3D printer of the same CraftBotXL make and model.Figure 20.

All the Kanawa-tsugi connections were printed on a 3D printer of the same CraftBotXL make and model.Figure 20. Load vs. displacement response for Kanawa-tsugi connectionFigure 21 shows the connection before and after testing. The failure occurred in the reduced cross-sectional area at the top of the connection, which is the weakest point. The failure occurred perpendicular to the layer orientation where the tensile stresses are largest. Figure 22 shows the failure mechanism of the connection, which can be considered a brittle failure since the decrease in load capacity is rapid, although not as sudden or brittle as the Kawai-tsugite connection.Figure 21. (a) Before and (b) after testing596Figure 22. Four-point bending test for Kanawa-tsugi connection from two different angles6.

(a) Before and (b) after testing596Figure 22. Four-point bending test for Kanawa-tsugi connection from two different angles6. Discussion of ResultsAfter testing three specimens of each connection type and three specimens of the reference beam with no connections, the average maximum load and deflection at the maximum load were calculated and are shown in Table 1. The ConXtech and Kawai-tsugite connections performed better that the Kanawai-tsugi connection, resisting a maximum load around 2450 N [550 lbs.] due to the better use of the total cross-sectional area compared to the Kanawai-tsugi connection. As expected, the connections had relatively large deflections at failure, which are about L/67 to L/47 (where L is the span between supports in inches).

As expected, the connections had relatively large deflections at failure, which are about L/67 to L/47 (where L is the span between supports in inches). For reference, for beams in building construction, the max deflection allowed for serviceability under service (i.e., not ultimate) loads is typically on the order of L/240 or L/360 [37]; thus, the deflections exhibited in the tested specimens, where the connections were placed at midspan, may not be capable of meeting serviceability limits in current building codes. However, a typical connection will likely be placed at the beam ends or at locations of lower moment demands, which would likely reduce the total beam deflection caused by connection rotations compared to the case where the connection is placed at midspan, the location of the maximum moment.

Some of the initial deflection is partially attributed to the tolerance gaps of 0.508 mm [0.02 in.] provided in the connections that had to “close” before the connection was fully engaged, after which the deflections are primarily attributed to material deformations.Table 1. Comparison of the average behavior of different connectionsReference Beam Kawai-tsugite Kanawa-tsugi ConXtechDeflection at Max. Load (mm) 3.10 4.37 2.95 3.51Max Load (N) 9,157 2,387 1,639 2,539Max Moment (kN-mm) 233 61 42 64It is worth noting that establishing an analytical expression to calculate the flexural strength of a 3D printed beam or connection is difficult due to the anisotropy associated with the printing direction and due to variations in print infill patterns. For example, uniaxial tests conducted on single strands of PLA filament, suggest that the material has a tensile strength of about 54 MPa [7,832 psi] [38].

For example, uniaxial tests conducted on single strands of PLA filament, suggest that the material has a tensile strength of about 54 MPa [7,832 psi] [38]. For a solid beam made of this material, it would be expected to have a flexural capacity of 332 kN-mm [2937 lb-in] per the ASTM D6272-17 standard [35] for a four-point bending test of the dimensions used in this study. The average flexural capacity of a reference beam (with 70% infill) in this test is 233 kN-mm [2,062 lb.-in], or approximately 70% of the theoretical strength of a solid beam made of the isotropic PLA material. Choosing a different infill pattern and infill percentage will result in different strengths. Future work could include specimens597with different infill patterns and infill percentages to investigate the impacts of these design parameters on beam and connection strength.

Future work could include specimens597with different infill patterns and infill percentages to investigate the impacts of these design parameters on beam and connection strength. When compared to the reference beam, the specimens with the Kawai-tsugite and ConXtech connections exhibited a flexural strength that was approximately 27% of the strength of a simple beam with the same infill pattern and percentage. As expected, the introduction of splice connections does decrease the capacity of a beam, but the proposed fastener-free connection is still able to transfer significant load and could be suitable for non-structural applications. It is worth noting that the American Institute of Steel Construction (AISC) [39] references Hart and Milek [40] who propose that “splices in fixed-ended [steel] beams be located at the one-sixth point of the span and be adequate to resist a moment equal to one-sixth of the flexural strength of the member, as a minimum”.

Also, Mohr & Murray [41] tested typical bolted steel beam splice connections, some of which exhibited flexural strengths of approximately 25% of the beam’s nominal flexural capacity. Thus, the additively manufactured fastener-free PLA beam splice connections tested in this study have relative strengths, compared to a non-spliced beam of the same material and dimensions, that are comparable to splice connections used in conventional steel construction. These fastener-free “quick connections” can therefore be considered as an alternative to conventional fastening methods for connecting additively manufactured parts, which could potential reduce the labor and time required to connect parts in the field.6. 3D Printing Limitations3D printing components introduces new design variables that can affect the strength or performance of an end product.

3D Printing Limitations3D printing components introduces new design variables that can affect the strength or performance of an end product. These design parameters for material extrusion include, but are not limited to, percentage of infill, the type of infill pattern, scale of the part being printed, the layer and print orientation, the layer thickness, the speed of the nozzle, temperature of the nozzle, temperature of the bed, temperature of the room, material use (e.g., PLA, ABS, nylon, etc.), 3D Printer model, and even the manufacturer of the material. To ensure the quality of 3D printed parts used in industries such as construction, more tests must be done to fully characterize the performance and variability of materials available in AM, to ensure environmental control and repeatability of results from a 3D printer, and to reliably predict properties of the end product based on the design parameters previously mentioned.

For example, this study only considered one infill percentage and pattern (70% infill in a square grid pattern), which was consistent between all specimens so that these tests could investigate the effects of connection type on strength and failure mode of each connection. Other infill percentages or patterns, such as a honeycomb or cubic pattern, could further improve specimen strength and/or ductility by ensuring the infill crushes in compression before rupturing in tension.One of the issues encountered with 3D printing components using common 3D printers is the cross-sectional area or the volume that can be attained by tabletop printers. For a 38 mm x 38 mm x 254 mm [1.5 in. x 1.5 in. x 10 in.] specimen, the printing time was around 19 hrs., which may be considered too long for on-site printing on a construction job site.

For a 38 mm x 38 mm x 254 mm [1.5 in. x 1.5 in. x 10 in.] specimen, the printing time was around 19 hrs., which may be considered too long for on-site printing on a construction job site. When considering large- scale construction applications, higher volumes of material will be used, and commonly available printers are not capable of producing components of such scale in the quick print time desired for most construction projects. Other institutions and companies, such as Big Area Additive Manufacturing (BAAM) from Oak Ridge National Laboratory (ORNL) [30] and Thermwood Large Scale Additive Manufacturing (LSAM) [42], are developing printers suitable for large-scale prints, and this is an ongoing area of research. Will there be a relationship between the performance598of a connection of 38 mm x 38 mm [1.5 in. x 1.5 in.] cross sectional area compared to a connection of 76 mm x 76 mm [3 in x 3 in.] or 305 mm x 205 mm [1ft.

x 1.5 in.] cross sectional area compared to a connection of 76 mm x 76 mm [3 in x 3 in.] or 305 mm x 205 mm [1ft. x 1ft.] cross sectional areas or larger? Further studies are needed to investigate scale effects in AM.Additionally, only one-layer orientation was included in this study, mainly being aligned to the testing orientation or sometimes referred to as 0°-layer orientation. Studying the behavior of a 45° or a 90° orientation and comparing the results may be able to corroborate previous research that states 45° is the best layer orientation for 3D printed components [36]. However, even if one print orientation is deemed to provide the largest strength, it may require supports during printing, resulting in a longer or less efficient print.

However, even if one print orientation is deemed to provide the largest strength, it may require supports during printing, resulting in a longer or less efficient print. For example, some applications, such as attachment of small architectural finished, may have minimal strength requirements (e.g., only supporting a relatively small self-weight) and a designer may favor printing efficiency and speed over structural performance. In other applications, such as load-bearing structural elements, part strength and quality may be of utmost concern. Effects of each variable on print performance and final state performance must be carefully considered and weighed based on the needs for each particular application.As discussed previously, layer thickness can play a key role in printing time, print quality, and final state strength, depending on the bonding between layers and the thermal strains being induced.

There is research that suggest smaller layer thicknesses will increase thermal strains and stresses that could reduce strength of 3D printed parts, which could be a major concern for structural applications [28]. Printing components as fast as possible to reduce printing time is sometimes required, but it affects quality of the final part and increases the likelihood of failure of the component due to skipping steps, warping, over extrusion or under extrusion of the material, etc. during printing.Another issue that can affect performance is the temperatures required for the nozzle and the heated print bed depending on the material used. For example, PLA requires temperatures around 215 °C for the nozzle and 55 °C for the bed, but some material manufacturers recommend 210 °C and 60 °C for the bed.

For example, PLA requires temperatures around 215 °C for the nozzle and 55 °C for the bed, but some material manufacturers recommend 210 °C and 60 °C for the bed. The 3D printing community has explored the variation of the temperatures that result in improved aesthetics and mechanical properties, but further work must still be done reliably characterize and qualify mechanical properties of printed parts based on such printing variables.Another thermal issue encountered while printing was the difference of temperature between the 3D printer and the surrounding room temperature. During winter, when the specimens from this study were printed, the thermostat in the building where the printers were stored was kept at a low temperature to conserve energy. During the time of printing from December 2017 to January 2018, the temperature in Austin had an average high of 16 °C [62 °F] and an average low of 5 °C [42 °F].

During the time of printing from December 2017 to January 2018, the temperature in Austin had an average high of 16 °C [62 °F] and an average low of 5 °C [42 °F]. Thus, it is likely the room temperature was somewhere between those values. These low temperatures affected the printing quality as the heated bed and the nozzle during printing should be around 60 °C [140 °F] and 210 °C [410 °F], respectively. The researchers were unable to print the parts at another time of year due to limited availability of Longhorn® Maker Studio printers for research purposes during the academic sessions. This difference in temperature caused a “temperature shock” in some of the prints causing them to fail or warp. For this reason, many in the 3D printing community have built personal enclosures for their 3D printers to control their surrounding temperatures and reduce the temperature sensitivity of these materials.

Even for5993D printing of concrete, which is being explored in the construction industry, some have provided a type of enclosure for their work to maintain a steady temperature and environmental control [43].The connections in this proof-of-concept exploratory study were only tested in flexure. To better characterize structural performance of these connections, further tests could be done to look at the behavior of these connections in compression, tension, and shear. For example, Kanawa- tsugi is known to have a better performance in tension compared to other wood joinery [6]. Additionally, further work can be done to optimize the geometries and dimensions of specific connections to minimize stress concentrations and improve structural performance under various loading conditions.As part of this study, the use of other materials and 3D printing processes were considered as a potential test variable.

The use of ABS using material extrusion process and rigid resins using vat photopolymerization were investigated for use in this study; however, limitations in printing sizes and AM technical issues prohibited their use in this study. This planned phase of the study was intended to compare the performance and strength for one specific connection, Kawai-tsugite, when printed with different materials and/or AM processes. The first attempt to print a specimen using material extrusion of an ABS material was done using the same settings as the previous PLA prints, where the nozzle and print bed temperatures were adjusted to values recommended for ABS. The component failed at about a third of the way through the print due to warping in one of the corners. It was suggested to add rafts to the component to decrease warping, but after another attempt, warping was even worse (shown in Figure 23).

It was suggested to add rafts to the component to decrease warping, but after another attempt, warping was even worse (shown in Figure 23). For the third time, the components were printed one by one to reduce the amount of wasted material in the event the print failed, but the print warped once again and failed. It was concluded that the amount of infill percentage and the size of the 3D printed component were some of the main reasons the print continued to fail, despite attempts to mitigate the issues. The failures are believed to be due to the printer taking almost an hour to print the first layer, resulting in differential temperatures and shrinkage in some areas of the first layer that caused it to detach from the print bed.Figure 23. Failed print using ABS600For vat photopolymerization a Forms 2 printer from Formlabs [44] was used to print the same Kawai-tsugite connection.

Failed print using ABS600For vat photopolymerization a Forms 2 printer from Formlabs [44] was used to print the same Kawai-tsugite connection. Advanced settings were needed to place the supports required for the print instead of using the one-click print feature in the program. Using the one-click feature required our component to be printed in an almost 45° angle; however, this print orientation was not consistent with the goal of testing the component with a layer orientation similar to the previous PLA 3D printed connections. During the first attempt to print the connection in a 0°-layer orientation, the print kept failing within an hour after starting. Figure 24 shows images from the slicer software that suggested insufficient supports to hold the specimen in place while printing. Later prints were attempted with increased support diameters, support density, and adding manual supports, but the print continued failing.

Later prints were attempted with increased support diameters, support density, and adding manual supports, but the print continued failing. No further attempts to print the connections with other processes or materials are planned due to time constraints. As such, exploration of connection behavior using different materials and processes is suggested as future research.Figure 24. SLA slicer softwareWhile the current study was focused on connection of structural and nonstructural AM components for applications in the construction industry, lessons learned from this study could be applied to other AM connection applications. Additionally, the connection concepts investigated using PLA in the current proof-of-concept study study can be applied to future research investigating use of other materials and AM processes more suitable for construction, such as using material extrusion for cementitious materials or powder bed fusion for metals.

Connections are meant to resist and transfer bending moments as well as shear and axial forces from one element to another. These tests have shown 3D printed connections are capable of transferring bending forces and can be comparable to the performance of some splice connections used in conventional steel construction, as previously mentioned. Future research in the area of extrusion of cementitious material is needed to find clever and reliable ways of adding reinforcement to the material, as concrete and cementitious materials themselves are brittle and performing poorly when subjected to shear and tensile stresses. For cementitious materials, new connection geometries that601take advantage of the material’s high compressive capacity and minimize locations of tensile stress concentrations should be further explored.Other materials such as fiber reinforced polymer or metals would be recommended for fastener-free connections due to their enhanced tensile strength and ductility.

These materials are well known in AM and are being explored for other structural applications at smaller scales [45]. Information about material properties are available from the companies that produce these materials for AM, but the behavior of 3D printed components using these materials are not readily available and would require further testing. The performance of these materials are temperature dependent and are typically printed as solid pieces for mechanical parts. All materials currently used in AM, including cementitious materials, metals, and fiber reinforced polymers, all require a controlled environment to build reliable components. It is expected that with further research and testing, performance of connections printed using materials such as metals and fiber reinforced polymer, as well as adequately reinforced cementitious materials, can perform similar to or better than the connections printed and tested in this thesis.7.

ConclusionThe main focus on this study is the investigation of different fastener-free connection alternatives that could be utilized either for formwork, scaffolding, architectural fixtures, and other non-structural or structural applications such as beam-to-beam, beam-to-column, or beam splice connections. A PLA material was used in this study due to its availability and ease of use with commercially available 3D printers; however, this material is not considered reliable for structural or non-structural applications since it is biodegradable and is not considered to have a high strength necessary for structural applications.

Investigation of different fastener-free connection concepts using this material is intended to provide some basic information on the relative strengths and properties expected of an anisotropic component due to the layer by layer processes using a material extrusion process and the expected failure modes of the different connections.The purpose of this study is to test this concept of 3D printed fastener-free connections at a small scale and with current AM technology to explore the benefits, limitations, and potential uses of such connections in structural and nonstructural applications. The extrusion AM processes used in this study are beneficial for generating the unique and customized geometries for the connections that would be difficult to generate using other formative or subtractive processes.

Three types of fastener-free connection concepts were explored in this study—the Kawai-tsugite and the Kanawa-tsugi connections inspired from Japanese woodworking, and the ConXtech inspired connections. Each connection was tested in flexure and was compared to a plain beam reference specimen with no connection and comparable print parameters. Results showed that the Kawai-tsugite and ConXtech connections exhibit higher flexural strength than the Kanawa-tsugi connection. These findings confirm that Japanese joinery could be used for new applications in construction using AM process. Such fastener-free connections, like Kawai-tsugite, do not necessarily need to serve a structural purpose, but could be reliable enough to be considered as a type of quick-connect connection used for temporary structures like formwork to reduce assembling and disassembling time.

Results from testing these additively manufactured connections have shown both the benefits and the limitations of 3D printing.602To encourage further applications of AM in construction, further research is necessary to develop and test new applications and improve and standardized testing procedures for components made using AM processes. The tests in this study were done using specific AM process settings, but variability of those settings such as layer orientation, infill percentage and pattern, tolerances, temperature changes, and scaling effects need to be considered. The exploration of connections that have been tested in this paper could be extended by using different materials available such as nylon, carbon fiber reinforced polymers, ABS, cementitious materials, and metals.

Additionally, different AM processes beyond material extrusion can be explored, such as powder bed fusion that could be used for nylon and metals, vat photopolymerization, or direct energy deposition also used with metal powder.8. References[1]ISO/ASTM, "52900-15, Standard Terminology for Additive Manufacturing - GeneralPrinciples - Terminology," ASTM International, West Conshohocken, 2015. [2] D. D. Camacho, P. Clayton, W. J. O'Brien, C. Seepersad, M. Juenger, R. Ferron and S. Salamone, "Applications of additive manufacturing in the construction industry - A forward-looking review," Automation in Construction, vol. 89, pp. 110-119, 2018.[3] Atom, "Construction Manager Magazine," 2013. [Online]. Available:http://www.constructionmanagermagazine.com/news/skanska-claims-industry-first-3d- printed-cladding/.[4] Biswas, K., Lind, R., Post, B., Jackson, R., Love, L., Green, J., Tryggestad, L., Obendorf, A., Eikevik, L., Guerguis, M..

2016.” Big Area Additive Manufacturing Applied To Buildings,” Thermal Performance of the Exterior Envelopes of Whole Buildings XIII International Conference, pp. 583-590, December 2016.[5] R. H. Thomas and M. J. Horman, "Fundamental principles of workforce management," Journal of Construction Engineering and Management, vol. 132, no. 1, pp. 97-104, 2006.[6] T. Sumiyoshi and G. Matsui, Wood Joints in Classical Japanese Architecture, KajimaInstitute Publishing Co., Ltd. , 1991.[7] M. D. Kijewski, "THINKING THROUGH MAKING: Material and Process Intelligence Revealed in Architecture Through Design and Fabrication," Halifax, 2017.[8] E. Brammeier, "Kawai Tsugite Joint," 18 March 2015.[Online]. Available:https://www.ethan-brammeier.com/kawai-tsugite-joint. [Accessed 5 January 2017].[9] R. Noe, "Japanese Woodworking Madness: A Three-Way Wood Joint," 13 October 2015. [Online]. Available: http://www.core77.com/posts/41624/Japanese-Woodworking- Madness-A-Three-Way-Wood-Joint.

Noe, "Japanese Woodworking Madness: A Three-Way Wood Joint," 13 October 2015. [Online]. Available: http://www.core77.com/posts/41624/Japanese-Woodworking- Madness-A-Three-Way-Wood-Joint. [Accessed 5 January 2018].[10] Buzz-Net, "After All Japanese! "Kawai Joint" that can be connected from three directions the topic abroad!," 25 December 2015. [Online]. Available: http://buzz-is netnews.com/kawai_tsugite. [Accessed 5 January 2018].[11] S. Kobayashi, "This Japanese joint called "Kawai tsugite"," 19 September 2014. [Online]. Available: https://www.instagram.com/p/tHohChREew/. [Accessed 05 January 2018].603[12] M. Wandel, "Clever 3-way joint (Kawai Tsugite) explained [Video file]," 9 October 2015. [Online]. Available: https://www.youtube.com/watch?v=Qv0VEN105nI. [Accessed August 2017].[13] ConXtech, "ConX Chassis Based Modular Building System," 2009. [Online]. Available: http://www.conxtech.com/wp-content/uploads/files/ConXtech-FAQ-0213.pdf.

[Accessed August 2017].[13] ConXtech, "ConX Chassis Based Modular Building System," 2009. [Online]. Available: http://www.conxtech.com/wp-content/uploads/files/ConXtech-FAQ-0213.pdf. [Accessed 28 August 2017].[14] K'NEX, "About K'NEX," [Online]. Available: https://www.knex.com/about-knex.[Accessed March 2018].[15] AutodeskInc.,"Fusion360,"2018.[Online].Available:https://www.autodesk.com/products/fusion-360/overview. [Accessed August 2017].[16] SketchUp, "SketchUp," 2018. [Online]. Available: https://www.sketchup.com. [Accessed28 March 2018].[17] SolidWorksCorporation,"SolidWorks,"2018.[Online].Available:https://www.solidworks.com. [Accessed March 2018].[18] AutodeskInc.,"Inventor,"2018.[Online].Available:https://www.autodesk.com/products/inventor/overview. [Accessed March 2018].[19] AutodeskInc.,"AutoCAD,"2018.[Online].Available:https://www.autodesk.com/solutions/3d-cad-software.

[Accessed March 2018].[19] AutodeskInc.,"AutoCAD,"2018.[Online].Available:https://www.autodesk.com/solutions/3d-cad-software. [Accessed March 2018].[20] AutodeskInc.,"3dsMax,"2018.[Online].Available:https://www.autodesk.com/products/3ds-max/overview. [Accessed March 2018].[221] I. Gibson, D. Rosen and B. Stucker, Additive Manufacturing Technologies, New York:Springer, 2010.[22] American Concrete Institute (ACI) 2010. Specification for Tolerances for ConcreteConstruction and Materials, ACI 117-10.[23] Ballast, D.K. 2007. Handbook of Construction Tolerances, 2nd Edition, Wiley. [24] American Institute of Steel Construction, Steel Construction Manual, 14th Ed. (Table 1-22), Chicago: AISC, 2011.[25] CraftUnique,Available: https://craftbotusa.com/category/craftbot-xl-3d-printer. [Accessed 3 September 2017]. Available: Replicator+,""CraftBotXL3DPrinter,"[Online].[26] MakerBot,[Online]. https://www.makerbot.com/replicator/.

[Accessed 3 September 2017]. Available: Replicator+,""CraftBotXL3DPrinter,"[Online].[26] MakerBot,[Online]. https://www.makerbot.com/replicator/. [Accessed 8 January 2018]."MakerBot[27] formlabs, "When to Use Different Layer Heights," 12 March 2018. [Online]. Available: https://support.formlabs.com/hc/en-us/articles/115000015870-When-to-Use- Different-Layer-Heights. [Accessed March 2018].[28] A. A. D'Amico, A. Peterson and A. Debaie, "Effect of layer thickness on irreversible thermal expansion and interlayer strength in fused deposition modeling," Rapid Prototyping Journal, vol. 23, no. 5, pp. 943-953, 2017.[29] Slic3r,3D http://slic3r.org/about. [Accessed March 2018]."Slic3r G-codegeneratorforprinters,"[Online]. Available:[30] L. J.

23, no. 5, pp. 943-953, 2017.[29] Slic3r,3D http://slic3r.org/about. [Accessed March 2018]."Slic3r G-codegeneratorforprinters,"[Online]. Available:[30] L. J. Love, "Additive Manufacturing of Molds for Fabircation of Insulated Concrete Block," Oak Ridge National Laboratory (ORNL), Manufacturing Demonstration Facility (MDF), 2018.604[31] The University of Texas at Austin, "Longhorn Makerstudios," [Online]. Available:https://makerstudios.engr.utexas.edu. [Accessed 29 August 2017].[32] B. M. Tymrak, M. Kreiger and J. M. Pearce, "Mechanical properties of components fabricated with open-source 3-D printers under realistic environmental conditions," Materials & Design, vol. 58, pp. 242-246, 2014. "PLA vs. ABS: What's[33] K. Giang,the difference?,"[Online]. Available: [Accessedhttps://www.3dhubs.com/knowledge-base/pla-vs-abs-whats-difference. January 2018].[34] B. Hesse, "ABS or PLA: Which 3D printing filament should you use?," 26 September 2015. [Online].

January 2018].[34] B. Hesse, "ABS or PLA: Which 3D printing filament should you use?," 26 September 2015. [Online]. Available: https://www.digitaltrends.com/cool-tech/abs-vs-pla-3d-printing- materials-comparison/. [Accessed March 2018].[35] ASTM International, "D6272-17 Standard Test Method for Flexural Properties of Unreinforced and Reinforced Plastics and Electrical Insulating Materials by Four-Point Bending," ASTM International, West Conshohocken, 2017.[36] MakerBlock, "12 Ways to Fight Warping and Curling," 23 June 2011. [Online]. Available:https://www.makerbot.com/media-center/2011/06/23/12-ways-to-fight-warping-and- curling. [Accessed March 2018].[37] American Society of Civil Engineers, ASCE/SEI 7-16 Minimum Design Loads andAssociated Criteria for Buildings and Other Structures, 2016.[38] T. Letcher and M.

[Accessed March 2018].[37] American Society of Civil Engineers, ASCE/SEI 7-16 Minimum Design Loads andAssociated Criteria for Buildings and Other Structures, 2016.[38] T. Letcher and M. Waytashek, "Material Property Testing of 3D-Printed Specimen in PLA on an Entry-level 3D Printer," in ASME 2014 International Mechanical Engineering Congress & Exposition, Montreal, 2014.[39] B. A. Mohr and T. M. Murray, "Bending Strength of Steel Bracket and Splice Plates,"Engineering Journal, vol. 45, pp. 97-106, 2008.[40] Alec, "Lewis Grand Hotel teams with Andrey Rudenko to develop world's first 3D printed hotel, planning 3D printed homes," 8 September 2015. [Online]. Available: http://www.3ders.org/articles/20150909-lewis-grand-hotel-andrey-rudenko-to- develop-worlds-first-3d-printed-hotel.html. [Accessed November 2017].[41] Formlabs, "Form 2," 2018. [Online].

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed Paper3D Printed Intelligently Graded Functional Stiffness Foam for Sturdier Multi Stiffness MaterialsB. A. Emery1,3, D. Revier1,2, V. Sarkar2, M. Nakura2, and J. I. Lipton1,2,31Mechanical Engineering, University of Washington, Stevens Way, Box 352600, Seattle, 98195, WA, USA. 2Computer Science and Engineering, University of Washington, 185 E Stevens Way NE, Seattle, 98195, WA, USA. 3Department of Mechanical & Industrial Engineering, Northeastern University, 334 Snell Engineering Center 360 Huntington Avenue, Boston, 02115, MA, USA.Abstract Foams are ubiquitous, being used in applications such as padding, insulation, and noise isolation. Bonding different density foams together produces undesired stress concentrations and boundary effects.

Bonding different density foams together produces undesired stress concentrations and boundary effects. Creating controlled gradients in foam properties has been a challenge for traditional and AM processes. Here we show how to use a form of material extrusion called Viscous Thread APrinting (VTP) to produce foams with multiple stiffnesses and continuous gradients between different stiffnesses. We do so by varying the path speed during extrusion to control the production of microstructures. We compare the process of producing discrete components and those with gradients, showing that those with gradients have higher strength in plane during tension, have no discontinuities in out of plane stiffness, and are less prone to forming cracks at the boundaries.

We demonstrate the process in thermoplastic polyurethane (TPU).1667IntroductionThe field of additive manufacturing (AM), often referred to as "3D printing", has seen broad adoption across various applications, ranging from bioprinting for medical purposes to mechanical applications optimizing static and dynamic properties [1–6]. One of the unique features of 3D printing is the ability to program spatially variant mechanical properties by varying the underlying microstructure of an object. This is often used for things like acoustic cloaking, elastic cloaking, etc [7–9]. The microstructure is traditionally fabricated through two primary methods: 1) the explicit definition of the geometry or 2) implicit definition through tool path algorithms, such as volume-filling infills. In traditional manufacturing spatially varying properties is achieved through discrete lamination of materials of differing material properties.

In traditional manufacturing spatially varying properties is achieved through discrete lamination of materials of differing material properties. Our work pioneers the use of Viscous Thread Printing (VTP) to produce continuously graded materials within individual layers without the use of explicit cell design or intricate pathing algorithms. VTP uses the dynamics of the printing process to control the microstructure of the object and thus the underlying mechanical properties as a function of space.Previous work on spatially variant mechanical material properties spans the use of topology optimization and expert design of cellular geometry. For explicit cell geometry, the printer's resolution must supersede the cellular structure’s unit cell, typically by an order of magnitude or more [6,10–12]. These geometries necessitate intricate cellular design through structures like octet truss unit cells or topology optimization [13,14].

These geometries necessitate intricate cellular design through structures like octet truss unit cells or topology optimization [13,14]. Yet, the cell size and design constraints often limit the applicability of explicitly defined cellular geometries in printed materials. For example, if a cell, such as an octet, is used, the unit cell must be constrained to a constant size in order to have a constant lattice spacing. This leaves the only control variable to be the diameter of the octet’s trusses. Alternatively, in topology optimization the geometry is less constrained, but the design space is significantly larger and the problem quickly becomes intractable for large objects with a high degree of variation.Our previous work with VTP created homogenous cellular structures with thermoplastic polyurethane (TPU) on standard Fused Filament Fabrication (FFF) printers.

Through VTP, we generated foam structures at resolutions like the printer's native resolution by leveraging viscous thread instability (VTI), a phenomenon well-observed in 1D and 2D [15–18]. This principle, commonly exemplified by the pattern formed when honey is poured onto toast, enabled the creation of 3D structures when stacked. In the previous work we were able to control macroscopic material properties via VTP slicer input parameters and demonstrated the ability to control effective Young’s modulus of the structure [19].Building on that groundwork, this paper extends the utility of VTP to produce foams with multiple stiffnesses and continuous gradients between them[1,19]. By dynamically manipulating print parameters during slicing, we present an innovative method of generating graded foam structures on a standard FFF 3D printer. With this approach, we produce foams of varying stiffness along the principal axes, decreased print artifacts and boundary defects.

With this approach, we produce foams of varying stiffness along the principal axes, decreased print artifacts and boundary defects. In turn, these structures have improved failure performance over previous state-of-the-art by having higher fracture stress, fracture strain, and toughness. This showcases the potential of VTP to revolutionize the fabrication of graded foams for a multitude of applications [20].1668SlicerThe key parameters to determine the coiling behavior of VTP are the dimensionless velocity variableand the dimensionless height variable,where F is the translation speed of the printhead, C is the exit speed of the material from the nozzle, H is the print height, D is the nozzle diameter and α is die swell, which is a material and process dependent constant that is determined empirically. Each of these variables can be controlled either explicitly or implicitly directly through G-code commands.

Each of these variables can be controlled either explicitly or implicitly directly through G-code commands. Together, thecombination determines the type of coiling behavior exhibited and spatial frequency ofthe coiling as well as the coil diameter respectively.There are three main regimes of coiling that are of interest for VTP foams: translating coiling, alternating coiling, and accumulation shown in Figure 3. Translating coiling results from a specific combination of Translating coiling is interesting because it allows for the greatest number of interconnections of a single thread producing a structure that is dense while maintaining porosity and uniformity. Alternating coiling results in coils that are further spaced apart, producing fewer interconnections and a less dense structure.

Alternating coiling results in coils that are further spaced apart, producing fewer interconnections and a less dense structure. Finally, accumulation is the closest to traditional 3D printing techniques with a low and can produce parts that are nearly 100% dense and extrusion widths many times greater than the nozzle diameter via over-extrusion; however, it does not result in porous structures.to produce a continuous path of overlapping loops of material.and lowControlling the extrusion speeds of the printing process. A G-code command G1 instructs the printer to move from its current positionof a printed object is done implicitly by setting the correct print and(𝑥, 𝑦 0) , 𝑧 0 0to a new position() 𝑥, 𝑦, 𝑧at a certain speed , while𝑓F , where length of material. In a G-code file this looks like G1 X Y Z E extruding G1 and the bold capital letters are alphabetical characters, and the lowercase letters represent numerical positions for the printer to parse.

In a G-code file this looks like G1 X Y Z E extruding G1 and the bold capital letters are alphabetical characters, and the lowercase letters represent numerical positions for the printer to parse. The infill toolpath, whereas∆𝑒𝑥 𝑦 𝑧 ∆𝑒 𝑓𝑥 and values to be solved for.and𝑦positions are determined by the object’sand𝑧, ∆𝑒𝑓1669The equations to generate G-code information frominputs are then𝐸˙ where G-code move, andis the constant speed to feed filament into the nozzle,𝐿is the Euclidean distance of the𝐴and𝐴are the filament and extruded thread areas respectively.𝑇 To generate spatially varying between them are called the transition regions.

For notation we employ a subscript/superscript distinction, where the superscript will always denote a mesh-wise quantity, while a subscript will always denote a point-wise quantity.𝐹we must first define areas of constantand regions1670First, we load multiple submeshes into our custom slicer where𝑗 Ωrepresents submesh j,∂Ω𝑗represents the surface of𝑗 , and ΩΩ = ⋃𝑗∈𝑀Ω𝑗is the total unioned mesh for the set of meshes M.These meshes are assumed to have constant cross-section in z and be non-overlapping but have some interfacing boundary. Each submesh is given a pair assigned by the user, which we combine in the notation spacing, effective layer height, transition length and toolpath discretization length, and die swell are also set by the user up front..

Other global variables, such as toolpathWe begin the slicing process by taking generates a toolpath traversing the entire space sections into subsection of lengthΩand feeding it to a traditional slicer (Slic3r) [21], which . We then refine each of the linear toolpathΩ𝑑≤𝑑which establishes the resolution of the𝑚𝑎𝑥transitions.Our transition algorithm associates each point along the toolpath with interpolated𝑝 𝑖π𝑖valuesby determining the weighted average of all the relevant submesh values. Relevant submeshes aredetermined by finding the distance between point𝑝 𝑖to∂Ω𝑗 . Here we use the signed distancebetweenand 𝑝 𝑖∂Ω𝑗 , where𝐷𝑗𝑖(, ∂Ω = 𝑆𝐷𝐹 𝑝 𝑖𝑗)and𝐷𝑗𝑖< 0⇒𝑝 𝑖𝑗 . A point is said to be in a ∈ Ωtransition region𝑇 Ωif the distance from point𝑝 𝑖to some∂Ω𝑗is less than the user definedtransition length, that is|||𝐷𝑗𝑖||| ≤ 𝐷𝑇⇒ 𝑝 𝑖∈ Ω𝑇.

Then the weights contributed by each submeshare:and thevalues at pointareMechanical Testing MethodsTo test our slicer algorithm we compared samples made using our algorithm using various transition lengths, with the base materials and a simple base slicer. For all specimens labeled TLXX we implement the above method for a bisected ASTM dogbone structure. For specimens labeled V0.15 and V0.40 we slice those with a normal slicer and alter the output G-code file directly using equations 3-5 through a custom python script. Finally, for the specimen labeled Zipped we slice each half of the dogbone separately and then “zip” the two separate G-code files together where layer 1 of the first file is printed and then layer 1 of the second file is printed in sequence, and so on.

This represents the previous state-of-the-art for multi-material printing of VTP foams [1].16711672For our validation we printed ASTM Standard D638-10 Type III dogbones out of NinjaTek NinjaFlex TPU [22]. The dogbone STL was split in half at the midway mark and various transition lengths were applied to create a linear gradient in dogbone. We conducted localized compression tests at 5 mm increments with a 10 mm diameter probe along the transition length to capture the variation in compressive modulus of the foams. The probe was printed on a Carbon M1 out of UMA-90 [23]. We then conducted ASTM Standard D638-10 tensile tests on an Instron 68SC-2 to measure the effective tensile modulus of the dogbone specimen as well capture fracture stress, fracture strain, and toughness.along the axial direction of theResults Figure 5 illustrates the impacts of the two slicing methods.

The upper row presents the result of our new slicing algorithm creating a continuous toolpath, while the lower row depicts a specimen created using the zipped method. Both cases reveal a distinct interface between. However, a noticeable 'bump' artifact emerges in the side view of the zipped version,resulting from the overlapping printed paths at the interface. The two slicers have qualitatively different edge crispness, as the zipped method introduces thickening at the edges. Additionally the non continuous movements of the zipped method introduces travel artifacts that are embedded in the print.This result illustrates the utility of our new slicing method for VTP foams in padding applications such as padding such as automotive seats, orthotics, and personal protective equipment. The defects from zipped sections would create unpleasant pressure points in these applications.

The defects from zipped sections would create unpleasant pressure points in these applications. Our new slicer by contrast produces structures with a less pronounced pressure point.Figure 4 details the local compressive modulus testing conducted on the dogbone structure. Here, linearly increases from 0.15 to 0.40, with varying transition lengths represented by vertical1673andblack lines for clarity. The compressive modulus remains largely unchanged outside the transition region, whereas inside this region, we observe a distinct modulus alteration. This phenomenon is further investigated in Figure 6 where an exponential curve of the formγ𝑥𝑓(𝑥) = λ𝑒 normalized to length one, allowing for direct comparison of the exponent term, corresponding to the line’s slope in the semi-log plot. As the transition region increases (i.e., acquires more data) the transition lengths greater than zero.is fitted solely to the transition regions.

As the transition region increases (i.e., acquires more data) the transition lengths greater than zero.is fitted solely to the transition regions. In each subplot, the transition region isγterm converges which indicates that the modulus as a function ofis constant forFigure 7 shows an abbreviated timelapse of the tensile tests from each TL family. The images capture three critical points: test initiation, mid-test at the 35-second mark, and the test conclusion. From these elements we can see that the transition length leads to distributed strain across the top and bottom halves of the samples for transition lengths 20mm and larger. This creates a more uniform transmission between the sections under loading.Figure 8 presents the mean stress-strain curve for each TL family, enclosed by their respective max and min boundaries. Due to variability in test durations within each TL family, data was truncated at the shortest test time to enable average calculation.

Due to variability in test durations within each TL family, data was truncated at the shortest test time to enable average calculation. The stiffest specimen (V015) is truncated vertically in the topmost plot to focus on the transition specimens, while the most compliant specimen (V040) lies beneath the transition specimens. The stress-strain curves for all TL families closely align, suggesting similar performance in modulus, fracture stress, fracture strain, and toughness. Table 1 and Figure 9 summarize the results from these stress-strain curves.1674Discussion In the discussion of our results, we found that dogbones with transitional regions exhibit similar elasticity to naively zipped dogbones. However, they notably outperform the zipped versions, exhibiting higher fracture stress, enhanced fracture strain, and superior toughness.

However, they notably outperform the zipped versions, exhibiting higher fracture stress, enhanced fracture strain, and superior toughness. This improved performance holds even in the context of zero-length transitions (TL00), where the VTP-sliced dogbones outperformed the zipped versions. This is attributed to the continuous toolpath created by our method.Table 1 shows the tensile modulus averaging 2.046 MPa across all transition families with a standard deviation of 0.115 MPa while the zipped dogbone has a modulus of 2.185 MPa. This indicates that tensile modulus of the dogbones is nominally constant as long as the average value ofacross the length of the dogbone is constant. We defineto be the integral ofover the entire length of the dogbone where desired transition length. In the case of a linear values multiplied by the dogbone length. This value is constant for both VTP and minimum sliced dogbones as well as the zipped dogbones.

In the case of a linear values multiplied by the dogbone length. This value is constant for both VTP and minimum sliced dogbones as well as the zipped dogbones. There is a clear indication that a simple method of determining dogbone tensile equivalence is to compare theis the location on the dogbone andis theit is equivalent to the average of the maximumvalue obtained from a curve.While Figure 4 and Figure 6. We average the values of the exponential fits ( determine a general fit for the transition regionis linear in, the compressive moduluschanges exponentially as seen inand ) from Figure 6 toThe modulus is also exponential in logarithmic definition ofbecauseis linear in . This means that with aa linear transition of modulus is achievable.16751676Table 1 also shows that the fracture stress, fracture strain, and toughness are consistent regardless of transition length, while the zipped dogbones performed at a much lower level.

Predictably, the zipped dogbones, without a continuous toolpath between the differing much earlier than the continuous transitions, likely due print errors and stress concentration at the interface. Compared directly to the TL00 sample, there was an approximate 33% increase in fracture strain and a 10% increase in fracture stress simply by using our continuous toolpath method without any transition length. This demonstrates that simply having a continuous toolpath between disparate areas can significantly improve the reliability of the sample and further benefits are seen when applying the transition over a region.regions, fracturedTable 1. Summary statistics for the tensile tests performed on the dogbones. Each row represents the average followed by standard deviation of 5 tests on dogbones printed with the same parameters. The mean of averages and mean of standard deviations of the samples printed with the new technique are given.

The mean of averages and mean of standard deviations of the samples printed with the new technique are given. The previous state-of-the-art and homogenous dogbone summaries are given at the bottom.

We presented a novel slicing approach that generates seamless toolpaths over several regions, each possessing distinct VTP characteristics. This technique enabled us to create smooth transitions between these regions, serving to minimize stress concentration and enhance the overall reliability of the interface. By applying our method, we observed a 20% increase in fracture stress, a 42% extension in strain before fracture, and an 84% improvement in material toughness compared to conventional methods that don't allow continuous transitions. We further detailed the compressive modulus along the transition, noting its non-linear relationship to modulus to be unaffected by transition length., and demonstrated the tensile1677It is clear that arbitrary control of the modulus in the transition region would be beneficial. Currently, we have determined a simple model to predict the compressive modulus as a function of space and transition length.

Currently, we have determined a simple model to predict the compressive modulus as a function of space and transition length. In future studies we intend to apply inverse design and determine the appropriate non-linear mapping of while this work focused on 2.5D structures, we would like to pursue graded material in all three dimensions to fully unlock the possibilities of graded material.to achieve a desired modulus layout. Additionally,References1.Lipton JI, Lipson H. 3D Printing Variable Stiffness Foams Using Viscous Thread Instability. Sci Rep. 2016;6: 29996. doi:10.1038/srep299962.Tappa K, Jammalamadaka U. Novel Biomaterials Used in Medical 3D Printing Techniques. J Funct Biomater. 2018;9. doi:10.3390/jfb90100173.Lee J-Y, An J, Chua CK. Fundamentals and applications of 3D printing for novel materials. Applied Materials Today. 2017;7: 120–133. doi:10.1016/j.apmt.2017.02.0044.Derby B. Printing and prototyping of tissues and scaffolds. Science. 2012;338: 921–926.

 Printed Smart Mold for Sand Casting:Monitoring Pre-Pour Binder Curing Nathaniel Bryant1, Janely Villela2, Juan Owen Villela2, Alan Alemán3, Josh O’Dell1, Sairam Ravi1, Jerry Thiel1, Eric MacDonald21 Metal Casting Center, The University of Northern Iowa, Waterloo, IA, USA 50614 2 The University of Texas at El Paso, TX USA 79902 3 Tec de Juárez, Ciudad Juárez, MéxicoAbstract: The benefits of additive manufacturing for fabricating complex sacrificial sand molds for geometrically-complex metal castings is revolutionizing the foundry industry driven by a digital manufacturing paradigm. The design freedom of 3D printing allows for new mold designs - not possible with traditional approaches - such as helical sprues, varying wall thickness to tailor the thermal history, and spatially-varying lattice castings. However, research on the curing time of printed molds, including the aging of printed molds, requires more exploration.

However, research on the curing time of printed molds, including the aging of printed molds, requires more exploration. This study describes the experimental evaluation of 3D printed specimens in which embedded environmental sensors were fully encapsulated into sand blocks during an interruption of the binder jetting process. Subsequently, over a 28 day duration, humidity, volatile organic compound generation, three environmental treatments. Mechanical testing of standard test specimens subjected to the same conditions was conducted. The sand structures held in high (uncontrolled) humidity and at reduced temperature were statistically significantly weaker than a third treatment based on the hypothesis that high humidity and/or low temperatures impede curing.

The use of embedded sensors could provide guidelines for mold and core storage conditions as well as in high-value production to inform the minimum (for full curing) and maximum duration (mold expiration) after printing to identify the optimal time to pour metal during the life of a printed sand mold.temperature and barometric pressure were captured forKeywords: Additive Manufacturing; 3D Printed Sand Casting; Binder Jetting; Curing;1.0 IntroductionTraditional sand casting is an economical metal manufacturing process that has been employed for thousands of years [1] and when compared to other casting processes (e.g. die casting, investment casting, etc.), sand casting can serve a diversity of alloys and can accommodate large casting structures. As the most common form of casting, sand casting produced over 100 million metric tons [2] of structures in 2020 [3].

As the most common form of casting, sand casting produced over 100 million metric tons [2] of structures in 2020 [3]. Additive manufacturing is now being explored to generate sand molds and cores with complex and customized geometries with binder jetting [4] - one of the seven sub processes within the additive manufacturing taxonomy [5]. A serendipitous benefit of additive manufacturing is access between the printing of layers to insert components inside of the mold and new strategies are now possible in the design, ventilation (channels to1528improve permeability), gating (sprues, runners to reduce the velocity of the molten metal to mitigate turbulence and reduce porosity [6]), and risers to manipulate the solidification.3D printing of sacrificial molds and cores involves binder jetting in which a binder is selectively inkjetted into a bed of sand premixed with an acid to provide the two-part Furan curing reaction.

After one layer of binder is jetted, the bed descends by a layer thickness (typically 250 microns) and a new layer of unbound sand is deposited uniformly across the bed and the process repeats. Interrupting the process and gaining access to the bed (either manually in this case, but potentially eventually with integrated robotics) is possible with ExOne SMAX printer and the thermal conduits, electronics with sensors, capability to integrate components (e.g. chills, actuation for dynamic gating systems, etc.), and subsequently, to continue the process to construct smart molds is now possible and relatively unexplored.

chills, actuation for dynamic gating systems, etc.), and subsequently, to continue the process to construct smart molds is now possible and relatively unexplored. Sand printing has been investigated at least preliminarily with explorations of the sand and binder materials [7–9], final casting characteristics such as porosity and surface finish [4,10,11], multiple-material castings [12,13], casting modeling [14], complex casting geometries [15–19] and even embedded sensors in molds and cores for unprecedented process monitoring [20,21] for detection of core shifts with sensitive and miniaturized accelerometers and gyroscopes [22], proper ventilation in molten-metal-engulfed cores [23] and even cyber security for the 3D print of sand molds [24].The present effort has investigated the insertion of fully embedded sensors into 3D printed sand structures during process interruptions in the binder jetting of Furan-based sand molds.

The driving hypothesis is that humidity and temperature (independent input variables) could be correlated to volatile organic compound production (dependent variable) and finally to the mold strength at the time of the metal pour. As potential future experiments may include extended durations (one year or more) of data collection and reporting, small Linux computers were employed that were not embedded in the structure to avoid the use of batteries as required by fully isolated sensor systems, thus allowing for virtually endless data collection. A recently developed and commercially-available four-mode sensor was inserted into a mold cavity that was exposed during a binder jetting interruption. A thin channel for a four-conductor cable allowed a Grove sensor connector to exit the structure but with minimal exposure of the internal sensor to the out-of-mold ambient conditions.

A thin channel for a four-conductor cable allowed a Grove sensor connector to exit the structure but with minimal exposure of the internal sensor to the out-of-mold ambient conditions. Pressure, temperature, humidity and volatile organic compounds (VOC) were measured: temperature and humidity were considered control variables that directly impacted the effectiveness of the curing. VOCs were considered to be a product of the curing process and an indication of the extent of curing and a predictor of final mold strength. Barometric pressure was not expected to be of utility but provided an indication of the local weather which impacted the independent variables in the uncontrolled ambient treatment.15292.0 Materials and MethodsAn industrial standard sand printer was used with traditional powder and binder consumables.

The process was interrupted and a portal was accessed to directly and manually manipulate the printed sand structures in mid-fabrication for the integration of sensors in the structure2.1 Printer and Binder Test specimens were produced using an ExOne S-Max 3D sand printer (Figure 1) utilizing an 80 AFS-GFN round grain silica sand and a furan binder system. The recipe employed a 65% virgin / 35% recycled sand mixture and an activator dosage level of 0.17% based on the batch weight. This dosage level was contingent on the observed flowability of the sand during the layering process.

This dosage level was contingent on the observed flowability of the sand during the layering process. The voxel dimensions used in this experiment were 0.1mm x 0.1016mm x 0.28mm, which resulted in a resin content of 1.2 - 1.25%, as determined by loss on ignition testing.Figure 1: ExOne Sand Printer at the University of Northern Iowa2.2 Test Specimens and Sand Block with Embedded Sensors Test specimens were produced using an ExOne S-Max 3D sand printer utilizing an 80 AFS-GFN In the center of the cube was a cavity that was round grain to create cubes 75 mm on a side. larger in dimensions than the BME680 Grove sensor board and a small channel was introduced1530from the internal cavity to the outside to allow for the exit of a 300 mm long Grove connector cable. The sensor location within the block was centered to maintain at least 75 mm of bound sound in all directions with the exception of the cable channel.

The sensor location within the block was centered to maintain at least 75 mm of bound sound in all directions with the exception of the cable channel. In addition, with each block of twelve sand blocks, several flexural test specimens were included in both the Y axis (in direction of the powder recoating) and X axis (in direction with the inkjetting) based on AFS method 3348-18-S. These specimens were treated in the same manner as the sensorized blocks for the 28 day period. Figure 2 illustrates the sand blocks during the process interruption ( left) and after completion (right).Figure 2: Sand block design at interruption (left) and after completion (right).2.3 Experimental Treatments Of the 12 sand blocks, sets of four (with two Raspberry PIs reading a pair of sensors each based on a limitation of the I2C bus addressing with a Grove Pi Hat used to connect the sensors) were subjected to three experimental treatments. Table 1 describes the conditions for each treatment.

Table 1 describes the conditions for each treatment. The hypothesis of the experiment is that temperature and humidity would impede binder curing and reduce mechanical strength. The ambient treatment was to simulate a typical foundry scenario with changes in weather. Furthermore, volatile organic compounds were hypothesized to be an indication of binder curing completeness as an output of the chemical reaction. Volatile organic compounds are released in the process and the integral of measured VOCs could correlate with final binder strength and used to ensure that molds and cores can withstand the head pressure for the specific alloy, design and height to ensure casting quality and yield.

The intent of the sensor monitoring was to ensure both that (1) molds have had sufficient time to cure (minimum time before pour after print) and (2) eventually to measure how long a mold can be stored in a warehouse or foundry before the binder begins to degrade and lose strength (maximum time between print and pour) for a given set of ambient conditions.1531Table 1: Experimental treatmentsGroup NumberConditionsPurpose1 (four samples)Laboratory ambient conditions to investigate a typical usage case of sand molds and cores.Ambient environment in a lab in Iowa during April of 2022 with varying conditions.2 (four samples)Refrigerated environmentConstant low humidity and low temperature (nom 2 °C).3 (four samples)Oven at 35 °CConstant low humidity and warm temperature (nom 35 °C)2.4 Sensor Insertion Methodology As the sand blocks were printed in the sand bed and at the point in which side walls of the internal cavity were complete and still accessible, the printer was interrupted and the loose sand in the cavities and channels were evacuated.

The sensors were inserted with cables and additional unbound sand was backfilled into the remaining volume of the cavities to provide a flush sand surface upon which printing could resume. The printing process was restarted to finish the top half of the sand cube and to entrap the sensor inside the structure while maintaining electrical connections through the external four-conductor cable. The main requirement was that only unbound sand be higher than the recoating level to avoid leaving voids in the sand bed..2.5 Standard Flexural Strength Testing All test specimens were tested for both the X and Y axis according to AFS method 3349-18-S for flexural strength. The resultant stress was calculated from the measured load using equation 1.σ = 3𝐹𝐿 2 𝑏𝑑Eq.

The resultant stress was calculated from the measured load using equation 1.σ = 3𝐹𝐿 2 𝑏𝑑Eq. 1Where F is the measured load required to break the specimen, L is the span length, and the denominator (bd2) is the cross-sectional area of the specimen.3.0 Results and DiscussionThe twelve blocks were binder jetted and a time lapse video of the sensor insertion process is shown here (Video 1). For a total of 28 days, data was collected. This effort and future work in long duration experiments are expected to provide guidelines for the storage of molds with minimum and maximum durations for a given set of humidity and temperature conditions. However, as these sensors are inexpensive, the utility of using these sensors even in production environments is possible and could help ensure the quality and yield of high-value castings.

However, as these sensors are inexpensive, the utility of using these sensors even in production environments is possible and could help ensure the quality and yield of high-value castings. Figure 3 shows how the sensors and cables are inserted through a printer portal in an evacuated cavity and channel.1532Figure 3: Insertion of sensors and cables into the blocks during printer interruption.3.1 Sensor insertion Of the twelve blocks, one block on the second layer after reinitiating the printing process had a sensor lift above the powder bed level causing an obstruction during the subsequent powder recoating. A large, gaping hole remained as shown in Video 1 and the half-built block and recoater collided. The cavity in the sand bed was manually backfilled with sand to allow for printing the top half of the block separately. After harvesting as shown in the video, the two detached sections of the bad block were manually bound together after inserting the sensor.

After harvesting as shown in the video, the two detached sections of the bad block were manually bound together after inserting the sensor. This manually reconstructed block was included in the oven treatment and did not appear to have significantly different VOC behavior. Figure 4 depicts the harvesting with sensors embedded and the cables exiting the sand blocks. Finally, Figure 5 illustrates all six PIs reading 12 sensors directly after harvesting and just prior to subjecting the sand blocks to one of three treatments.Figure 4: Harvesting of sand blocks with sensor cables shown.1533Figure 5: 12 Sand blocks and sensors connected to six Raspberry PIs prior to being subjected to the three environmental treatments.3.2 Sensor data and the impact of local weather on curing For 28 days, humidity, temperature, volatile organic compounds and barometric pressure were collected every minute.

Although barometric data was not expected to provide utility in this investigation, one interesting discovery was that insights into the local weather and the potential impact on the binder curing of the blocks were uncovered. On day two of the duration, a snow storm descended upon Iowa and was manifested by a precipitous drop in the barometric pressure. During the period of reduced pressure, the humidity was shown to spike given the weather change and increased moisture in the air appears to have interfered with the production of volatile organic compounds.

Of the experimental treatments, the case of ambient conditions would include the effect of this one-day duration of increased humidity as shown in Figure 6.Figure 6: Barometric pressure (left), humidity (center) and VOC generation (right) of one ambient sensor is shown in which a snow storm was responsible for a local increase in humidity which is hypothesized to have reduced the VOC generation and thus impeded the curing process for approximately a day.15343.3 Sensor data for entire duration Of the three treatments, four sensors were read from a pair of Raspberry Pi 4’s each. However, for two of the Pis (one in the refrigerated treatment and another in the ambient treatment), sensor communication stopped even though WIFI communication with the PIs continued. An investigation for this failure is underway and the reason is unclear.

An investigation for this failure is underway and the reason is unclear. Both sensors associated with each of the PIs was not readable so it appears to be related to the PIs; however, all six PIs with 12 sensors were tested for over a month prior to the sensors being embedded in sand and all collected data continuously and correctly. The only difference was that the sensors were subsequently embedded in acid-soaked sand but in the case of both non-functioning PIs, the two sensors were unreadable leading the investigators to believe that the sensors were not the problem (i.e. if related to the sensor functionality, only one would stop functioning).

if related to the sensor functionality, only one would stop functioning). One possible explanation is that the sensors drew more power based on a resistive short between nodes on the board due to the moist sand and the increased power consumption caused a voltage droop in the PIs and affected the PI functionality.Figure 7 illustrates the temperature, humidity and VOC generation for the four sensors in the oven treatment (low humidity and higher temperature). As expected, the temperature was measured between 34 and 38 °C and the humidity was maintained below 25% for the entire duration and all four sensors provided consistent data. The VOC generation was maintained at a level of between 20,000 and 40,000 ohms which served as a proxy for the actual VOCs and provided a relative level which was generally consistent throughout the duration with a slight increase over time.

In all four sensors, two spikes are seen at approximately day 5 and 16 which correlated with unexpected power outages. The PIs were programmed to restart data collection if rebooted and the VOC sensor requires warming up with initial values that artificially spike for the first 30 seconds. The spikes provided additional information as an actual spike in VOC generation would generally last longer and the artificial spikes inform the user of a power outage or related reboot.The VOC trends for all four sensors are consistent but the absolute values vary widely as these sensors were not calibrated. Consequently, without a calibration, these values cannot be used other than identify trends during the duration within a single mold and are not suitable for the molds were sufficiently cured.

Consequently, without a calibration, these values cannot be used other than identify trends during the duration within a single mold and are not suitable for the molds were sufficiently cured. providing a final single metric value to indicate that Alternatively, the temperature and humidity values were consistent in terms of absolute value and together could be combined to provide a “pour / no pour” metric. The VOC sensors could be calibrated prior to use by measuring VOCs in a known VOC environment for a specific time. However, the temperature and humidity sensors are calibrated sufficiently out-of-the-box and could be used to calculate a single combined metric to determine the extent of the binder curing.1535Figure 7: Oven treatment sensor data for four sensors read by two PIs.The two sensors monitored in the refrigerated treatment are shown in Figure 8 with temperature and humidity both as expected.

The initial temperature is higher as the sensors were connected in the laboratory and later transported to the refrigerator causing an initial higher temperature. Only two sensors were included in this treatment as one of PIs stopped reading the additional two sensors. Similarly to the oven treatment, the conditions were consistent for the cold, low-but- slightly higher humidity case than the oven treatment. The VOCs in both cases appear to have a slow declining trend in value and although the shape of the graphs are very similar and show the power outage that was seen in the oven treatment data, the overall magnitudes of the graphs differ by a factor of three by the end of the experiment, further illustrating how the absolute values for this sensor is not valid but rather that the general trend for any given sensor may In this case, different from the oven treatment, the VOCs appear to slowly provide insights.

decline over the 28 day period.1536Figure 8: Refrigerated treatment sensor data for two sensors.The final ambient treatment is included in Figure 9 and also only includes two sensors for likely the same problem that affected the non-functional PI in the refrigerated case. The temperature and humidity vary more dramatically than the oven or refrigerated cases as expected in an uncontrolled environment and the VOCs in both sensors appear to remain relatively consistent throughout the duration. However, again, the absolute values are not consistent due to the lack of calibration.Figure 9: Ambient treatment sensor data for two sensors.3.4 Flexural strength of specimens In all eight successfully-read sensors, the values during the 28-day period of humidity and the binder curing which is temperatures were averaged for determining the extent of1537hypothesized to be correlated with the final mold strength.

Figure 9 shows the X and Y flexural strengths where X is the direction of inkjetting and Y is the direction of recoating. For both axes, the heated and low humidity case provided improved average strength and the average temperature and humidity over the duration of the experiment had strong correlation but not with statistical significance as the four of the sensors failed to report.. This correlation could be used to provide a threshold to identify the minimum time for curing to provide sufficient strength prior to pouring metal.

This correlation could be used to provide a threshold to identify the minimum time for curing to provide sufficient strength prior to pouring metal. Future experiments will include measuring conditions for months or years to establish a potential maximum duration for which the molds could be stored prior to use assuming the binder would eventually degrade over time.Figure 10: Flexural strength in the X and Y build directions for each of the three treatments: Heated, Refrigerated and Ambient.4.0 ConclusionsThe present experiment has demonstrated the basic utility of monitoring environmental conditions of printed sand molds and cores to ensure that binder was sufficiently cured. This technique will be further explored to establish guidelines for a window in time for which the1538molds could be used for casting depending on humidity and temperature.

conclusion points include:The importantSensors can be embedded successfully within molds during a printer interruption of sandmolds to inform the extent of curing,For the specific BME680 sensor, the Volatile Organic Compound sensor mode would require calibration to be used in determining if sand binder had sufficiently cured to provide the mold strength required for metal pouring,A combined metric of temperature and humidity has been preliminarily shown to provide a higher correlation between the calculated metric and the final mold flexural strength, ● With further experiments, guidelines could be established to allow for the monitoring of temperature and humidity in a mold storage facility to calculate a minimum curing time for any specific conditions, andLong-duration experiments will be the focus of future work to understand how long 3D printed sand molds remain sufficiently strong, and if these molds need to have an expiration date with the specific consideration of the environmental conditions that the molds were subjected to in the storage facility over time.AcknowledgmentsWe would like to thank the Murchison Endowment at the University of Texas at El Paso and a grant from the American Foundry Society.References1.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed Paper3D PRINTING ENABLED-REDISTRIBUTED MANUFACTURING OF MEDICAL DEVICES J. Munguia*, T. Honey1, Y. Zhang1 M. Drinnan2, C. Di Maria2, A. Bray2, M.Withaker21Newcastle University, School of Mechanical & Systems Engineering. Newcastle upon Tyne, UK 2 Regional Medical Physical Department (RMPD), NuTHC.AbstractRecently the home-use segment of medical devices has entered in the loop of Additive Manufacturing (AM) enabled optimizations, this includes CPAP masks, insulin delivery packs and diagnostic tools such as urine-flow meters. Here we analyze the supply chain provision of a specific uroflowmetry device which is originally designed in Europe, manufactured in Asia and which has a range of distribution channels across healthcare systems.

This paper analyses the impact of various AM technologies that can enable near-patient manufacture of devices on- demand. Our analysis shows that the cost of design-changes (or product updates), when reflected on the overall lifecycle cost, can be comparable to producing the device locally with a different supply chain arrangement. Furthermore it is suggested that in order to fully exploit the capabilities afforded by AM, the original product’s design features must be modified so that built-times are reduced allowing a larger 3D printing-based production capacity.Keywords: 3D printing, home-use medical device, redistributed manufacturingIntroductionCurrently home-use medical devices are manufactured in large numbers with the overall process being distributed across different countries. Often, devices are manufactured in one country, assembled in a second country, and then delivered to the end-user along a chain of distributors and selling points.

Often, devices are manufactured in one country, assembled in a second country, and then delivered to the end-user along a chain of distributors and selling points. The investment in tooling for mass- market devices is substantial and beyond the means of most start-ups and small- and medium-sized enterprises (SMEs) that represent more than 80% of the UK medical industry (BIS, 2010). Manufacture usually represents only part of the final cost of the device, with another large part being represented by costs related to shipping and storage at the industry site and then at the various distribution and selling points.The possibilities for redistributing manufacturing closer to the end-user (the patient) could have a positive impact on the final cost of the medical device, reduce the time for the device to be available to the patient, improve the patient experience, and offer an opportunity for customisation.

In this project, we explored the technical feasibility and the opportunities offered by Redistributed Manufacturing (RDM) applied to home-use medical devices. The aims were to:Investigate the technical feasibility of manufacturing an existing medical device using an RDM approach enabled by Additive Manufacturing.Quantify the cost of deploying the medical device using an RDM approach. • Evaluate the likely patient take-up of RDM medical devices.Methodology We considered an existing home-use medical device as a case study for this project (The Flowtaker, MMS/Laborie website), a disposable urine flowmeter for single-patient use which was developed by the Northern Medical Physical and Clinical Engineering Department (The Newcastle upon Tyne Hospitals NHS Foundation Trust, UK) and licensed to the major multinational manufacturer of instrumentation for urodynamic (MMS/Laborie).

In first instance, we developed a benchmark comparison between the original device and a modified version suited for Additive Manufacturing (AM) in terms of design and manufacturability features. This was done by characterizing the number of individual components that comprise the original device, defining the standard Bill of Materials and analysing particular manufacturability features of the original and new flowmeter.1680Next, a similar device with the same shape and geometric attributes was manufactured using Additive Manufacturing (Fused Deposition Modelling, FDM) and the design fully optimized to reach the maximum possible productivity for the selected FDM process. The final part of the study compared the supply chain models for both the conventional manufacturing route, and the envisaged redistributed manufacturing route that would be followed if AM were implemented.Results 1.

Design and manufacturability benchmarkIn order to describe the main manufacturability features and the subsequent logistic needs, it is necessary to understand the device’s design; the main body of the flowmeter was designed for injection moulding including typical features (Figure 1), such as stress relief lines, pins and runners space, gussets to avoid warping on thin walls, bosses and ribs for counteracting thermal shrinkage inside the mould and ribs to support the thin shell body. All this requires the design and manufacture of production tooling and metal injection moulds.

All this requires the design and manufacture of production tooling and metal injection moulds. Due to economies of scale and existing manufacturing expertise within the developing team original production moved to Asia particularly due to the need for assembly or electric/electronic components to be fitted on the plastic body.Figure 1 – Top and reversed-internal view of flowmeterThe current overall design has a part count > 30 components (Figure 2), including plastic elements, electronic components, wiring, connectors and instruction labels. One of the driving cost factors is the initial tooling investment as the device requires four individual injection molds to be sourced. With this in mind, the next step in our analysis was to replicate the plastic components and assessing the current capabilities of existing AM technologies.Figure 2 – Disassembled overview on main individual components1681Discussion.

Comparing 3D-printing methodsThe original engineering design files of the Flowtaker components were transferred to a 3D-printing (3DP) compatible format allowing us to replicate the same geometry using an additive process. In order to choose the most suitable 3DP technique, we accounted for the following factors: • Ease of operation in a medical setting. • Space requirements for equipment, material handling and post processing. • Cost of ownership or initial investment.Selective Laser Sintering (SLS) is one of the most established industrial-grade 3DP processes. It offers high printing speeds, minimising the fabrication time. SLS offers fine accuracy (typically 0.1mm) - though far from injection moulding - and overall good repeatability. However, the use of a laser would represent a safety issue in a hospital environment.

SLS offers fine accuracy (typically 0.1mm) - though far from injection moulding - and overall good repeatability. However, the use of a laser would represent a safety issue in a hospital environment. Additionally, the materials for SLS come in powder form with a particle size of 50-54 microns, which represents another safety risk and would require suitable breathing equipment and ventilation. Initial cost of ownership starts at $200,000 (Wholers, 2013).Stereolithography (SLA) is another widely accepted 3DP process that uses liquid photopolymers as building materials, which need ultraviolet light as the curing source. It also requires post-curing, which increases the total fabrication time. SLA usually requires the use of supporting structures to prevent deflection of the final product due to gravity, which involves post-processing to remove the supporting structures and it can be difficult to carry out for the lay person.

In addition, a limited range of photopolymers can be used as building materials and the materials cost ($100/litre) makes it less attractive than other processes. The printing quality of SLA is better with lower printing speeds; therefore, the printing quality and time length need to be balanced to build a reliable product using the shortest building time. Cost of ownership starts at $2,000 for low-cost units (Wholers, 2013) but can go beyond $150,000 for multi-capacity stations.Fused Deposition Modelling (FDM) was our selected method as it satisfied the three factors described above. FDM units can be operated in a medical environment with minimum noise and non-toxic fumes produced. The cost of ownership for desktop units starts under $1,000 (Wholers, 2013). However, the main advantage of FDM over other 3DP methods is its capability to use similar materials to those used for injection moulding. This can allow for manufacturing designs without the need for moulds and tooling.

This can allow for manufacturing designs without the need for moulds and tooling. FDM as an alternative process. Initial trials showed that, although it was possible to replicate the exact design of the original Flowtaker using FDM with plastic materials, the overall accuracy and part quality was far from the injection moulded version (Figure 3). Furthermore, it required up to 48h to print one single Flowtaker case.Figure 3. Comparison of FDM printed (a) and injection moulding (b) plastic part.Consequently, we decided to revise the design of the device in order to overcome the manufacturing constraints posed by the original design and to harness the advantages offered by FDM. This also required a complete re- design of the electronics board and sensing components. Figure 4 shows the new design of the electronic components and the corresponding polymer case with a printing time drastically reduced to 5h/case.

Figure 4 shows the new design of the electronic components and the corresponding polymer case with a printing time drastically reduced to 5h/case. The new design successfully consolidated and/or eliminated extra components, largely reducing manufacturing time and need for assembly.1682Figure 4. Optimized FDM concept2. Supply chain modelsAlthough the re-designed flowmeter is still far from the market, we envisaged a number of scenarios that could potentially benefit from an Additive Manufacturing-enabled redistributed supply chain model. The traditional in- clinic pathway to assess patients with urinary problems requires the patient to attend the hospital. The Flowtaker offers a novel possible pathway where the patient can perform the measurement in the comfort of his home as opposed to performing the tests at a designated hospital by appointment only.

The Flowtaker offers a novel possible pathway where the patient can perform the measurement in the comfort of his home as opposed to performing the tests at a designated hospital by appointment only. In the current manufacturing model, the manufacturing and assembling process is distributed across different countries; the final distribution to patients across the UK, Europe and North America takes place through a chain of various distributors. Manufacture only accounts for a fraction of the final cost whereas distribution, storage and handling take a considerable portion of the final retail price (Figure 5 and 6).Figure 5 Overall supply and value chain for the existing Flowmeter model1683Figure 6.

Simplified percentage contribution of various supply chain elements on the final product cost These figures clarify the role of ‘pure manufacturing’ as a minor contributor to cost and show that in a re- distributed manufacturing scenario the intermediate and later storage + assembly + shipping stages have a larger contribution. This is particularly relevant when exploring a 3D printing scenario as manufacturing costs when using additive technologies are normally higher when compared to traditional high volume production.We believe that the patient pathway offered by the Flowtaker could be further enhanced by the possibility to manufacture and assemble the device closer to the patient by using RDM technologies. We are exploring two different possible strategies: manufacturing close-to-patient, and manufacturing in-the-home. Manufacturing close-to-patient.

We are exploring two different possible strategies: manufacturing close-to-patient, and manufacturing in-the-home. Manufacturing close-to-patient. This model can open new opportunities for rearranging the conventional supply chain for home-use medical devices by gaining efficiency and reducing the overall lead times. In particular, we envisage a small community manufacturing facility in healthcare centres and pharmacies. In this scenario manufacturing takes place near-the patient environment and in non-dedicated/ non-industrial facilities which can be pharmacies (i.e. Boots) GP practice and regional hospital units.

In this scenario manufacturing takes place near-the patient environment and in non-dedicated/ non-industrial facilities which can be pharmacies (i.e. Boots) GP practice and regional hospital units. This local manufacture fashion is enabled by the adoption of office-friendly 3DPrinting units (most likely FFF technology) to be operated by a resident expert with enough basic knowledge to: • Receive hospital/GP prescription for a flow taker device • Prepare a digital 3D file according to prescription • Select appropriate 3D printing filament and prepare 3D printer parameters • Remove, post process and assemble 3D printed structure to off-the shelf electronics subassembly. In this re-distributed model (Figure 7) centralized overseas manufacture takes place only for internal circuit- board/electronics, then this is shipped to individual regional locations (pharmacy) where on-site 3D low cost 3D printers are used to produce the rest of the plastic components on demand.

In this new model we acknowledge the cost contribution of the initial design stage as it is assumed that the potential of 3D printing is based on personalization, design changes, upgrades and continued improvement, so this initial stage is set to remain active throughout the life of the new device supply chain.Figure 7. Re-distributed Scenario 1 and cost distribution Manufacturing at home. Although this scenario is currently technically feasible, it poses a challenge to the MedTech manufacturing industry to secure the quality and reliability of the medical devices which are printed at home, as all non-electronics manufacturing and assembly would be “outsourced” to the patient’s home. This model relies on the end customer having a 3D printing unit available either at home or at community centres (Fablabs, Makerspaces, online bureaus, etc.), so the cost of producing and assembling the device is placed on the1684back end on the supply chain.

We have estimated that the initial design stage would also increase its contribution to the final cost (Figure 8) as this stage will be key to design and generate 3D files which can be reliably exported and printed on a variety of 3D printing units. This will include 3D file integrity checks, 3D printing trials and materials testing before the medical device files are released to the public. This cost contribution also requires the manufacturer to setup and host a 3D file repository where users can download verified 3D print files for the relevant device in similar fashion to current online 3D printing repositories (thinginverse, grabCAD, youimagine, etc.)Figure 8. Re-distributed Scenario 2 and cost distributionThe final pricing of the device in this scenario remains a challenge as the patient’s contribution to manufacturing and assembly eliminates the intermediate storage, distributions and subassembly stages, potentially lowering the final price of the device.

It is more likely that this model will not completely substitute the conventional supply chain but it can be made available to a pool of customers who value the quicker availability at the expense of performing the assembly and non-critical parts manufacture in-house.Figure 9. An RDM supply chain scenario where manufacturing happens after the device has been prescribed to the patient.Discussion The new provision chain will need to be designed according to the service needed (Slack et. al, 2010) that means Redistributing Manufacturing will take different forms according to the regional customer needs. For home-use medical devices, this can mean production will be relocated closer to the points of higher prevalence of certain medical conditions, as this can critically reduce waiting times for treatment.

This is not currently possible by following conventional process routes; therefore, the combination of AM-enabled production and RDM principles can help plan and justify the creation of clustering locations where strategic manufacturing must happen.Quality control is currently performed at the centralized manufacturing locations and quality assurance carefully follows strict medical device directives and standards (MDD, 1993; BS-EN-60601, 2006). Whether it is dimensional accuracy, product/shape conformity or durability tests, a scenario where manufacturing is redistributed away from a central location can pose regulatory challenges: will regional hospitals become specialized manufacturers for niche devices? Will regulatory bodies open up to non-centralized production models?1685The current cost model comprises of the following elements: R&D, design, licensing, manufacturing, assembly, shipping and storage.

A novel AM-enabled cost model that considers variations in manufacturing scenarios can potentially eliminate both the shipping and storage elements. However, Baumers et al (2016) also identified additional sources of cost in AM-enabled fabrication such as post-processing and multiple-part building. AM- enabled manufacturing would also offer an opportunity for customization of the final device with respect to size, geometric specifications, functional features, colours and material.Current manufacturing can reach production capacities of several thousand devices per week. Based on our re- designed optimized device with a total print time of 3-5h, it would be possible to manufacture between 5–8 devices per day locally using a single FDM unit. Obtaining consistent throughput information will be vital for the control of the overall process.

Obtaining consistent throughput information will be vital for the control of the overall process. During our preliminary analysis we concluded that one AM process (FDM) has the potential to substitute the conventional process followed for our specific case study. However, it has been demonstrated that different AM technologies are also suitable for different medical devices such as foot orthotics SLS), CPAP respirators (DLP, inkjet) or hearing aids (SLA).Conclusion This feasibility study explored the possibilities for replacing existing operations normally adopted for a sample medical device case study. The flowmeter introduced as the basis of this study was redesigned and produced using alternative manufacturing methods and the main implications of such change have been discussed.

The flowmeter introduced as the basis of this study was redesigned and produced using alternative manufacturing methods and the main implications of such change have been discussed. Given the novelty of the topic and subject area it is not expected to perform a comprehensive analysis of the total supply chain per se, but to point at the immediate critical changes that a redistributed concept can bring into the medical device manufacturing arena.Multiple challenges for operations management were identified, the most crucial being regulatory, quality and costing issues. The advantages of flipping production and distribution from a centralized model to a strategically re-distributed model are still to be confirmed and two key factors will be: 1) patient take up, and 2) practitioners/healthcare systems adoption.

Given that the technical possibilities of such change are readily available, more feasibility studies taking into consideration different devices will be instrumental in documenting this potential paradigm change.AcknowledgementsThis project is supported by the UK- EPSRC Redistributed Manufacturing in Healthcare Network (RiHN), grant EP/M017559/1.ReferencesBIS (2010) Department of Business, Innovation and Skills, UK Government. Strength and opportunity. The landscape of the medical technology, medical biotechnology and industrial biotechnology sectors in the UK. MMS (2016) Medical Measurement Systems International. Flowmaster - PC based Wireless. Uroflowmeter http://www.mmsinternational.com/int/694/mms-products Wholers T, (2013) Wholers Report 2013, Additive Manufacturing and 3D P rinting state of the industry, annual worldwide progress report. Forth Colling, US. MHRA (2013) Medicines & Healthcare products Regulatory Agency. Custom-made devices.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference3D PRINTING IN THE WILD: A PRELIMINARY INVESTIGATION OF AIR QUALITY IN COLLEGE MAKER SPACESBill McDonnell, Xavier Jimenez Guzman, Matthew Dolack, Timothy W. Simpson, and John M. Cimbala Mechanical & Nuclear Engineering, Penn State University, University Park, PA 16802AbstractAdditive manufacturing is a popular method for prototyping and manufacturing custom parts, especially on college campuses. While there is widespread use of 3D printers as part of many engineering classwork, there is little regulation or knowledge regarding emissions. Many plastics, including polycarbonates, ABS, and PLA are known to emit high counts of volatile organic compounds (VOCs) and particulate matters (PMs).

Many plastics, including polycarbonates, ABS, and PLA are known to emit high counts of volatile organic compounds (VOCs) and particulate matters (PMs). This study focuses on VOC and PM counts in several natural environments and dedicated “maker spaces” on a large college campus to gauge the exposure that students and operators experience. Emissions were measured using a photoionization detector and two particle sizers. The photoionization detector measured total VOCs, and the particle size counters measured both total nanoparticles and individual micro- particles based on relative particle diameter. Measurements were taken in hourly increments and then analyzed to determine the degree with which desktop printers emitted VOCs and PM.

Measurements were taken in hourly increments and then analyzed to determine the degree with which desktop printers emitted VOCs and PM. Our data can be used to determine whether additional ventilation or filtration is needed when 3D printing “in the wild” to enhance operator and bystander safety.IntroductionAs the additive manufacturing industry has developed over the past several decades, the use of consumer-grade (desktop) 3D printers has increased as well. It is now very common for universities to augment their engineering and design programs to include instruction and use of 3D printers given their ability to help visualize ideas and solve complex design problems. Large universities like Penn State University offer a wide array of 3D printing capabilities for in class projects as well as research (e.g., materials development and characterization, multi-extruder systems).

These can range from low-cost, open source, do-it-yourself (DIY) 3D printers that are built as part of a class and used in other classes to higher-end polymer systems that are used for senior capstone design projects and fabricating prototypes for research. Many students even have desktop printers in their dorms. With the ever increasing use and popularity of desktop 3D printers, the issue of operator and bystander safety has come into question.Material extrusion technology, particularly fused deposition modeling (FDM) is rapidlyevolving, and as a result the technology is largely unregulated. This is concerning because studies conducted by universities and regulatory agencies have shown that FDM printing emits particulate matter (PM) and volatile organic compounds (VOCs) [1]. Additionally, studies have linked PM and VOC emissions, especially when in high concentrations, to numerous health problems [2].

Additionally, studies have linked PM and VOC emissions, especially when in high concentrations, to numerous health problems [2]. The emissions of greatest concern are VOCs, ultrafine particulates (UFP), and fine particles. Ultrafine particles and fine particles differ from each other in that ultrafine particles2456have a diameter less than 100 nanometers and fine particles have a diameter less than 2.5 micrometers [3]. Ultrafine particles are a subgroup within fine particles. Inhalation of fine particles, UFPs, and certain VOCs are scientifically proven to be carcinogenic [4].In total, approximately ten studies have been conducted by regulatory agencies, academicinstitutions and FDM printer manufacturers that have looked into FDM printer emissions. The studies mostly focus on VOC and UFP particle emission, and they typically use a flux chamber to precisely measure emissions.

The studies mostly focus on VOC and UFP particle emission, and they typically use a flux chamber to precisely measure emissions. One experiment used a flux chamber to measure emission rates with a TSI 3007 condensation particle counter and TSI 3910 Nano Scan scanning mobility particle sizer (SMPS) [5]. Another study measured particle concentrations in a modified office environment again using a TSI 3910 Nano Scan SMPS [6]. Finally, one study used a sealed, cubic testing chamber with a HEPA filter to measure TVOC concentrations [2].With this in mind, we set out to study particle emissions and VOC concentrations in naturally environment at seven locations around Penn State’s University Park campus. The selected locations all contained FDM printers that are used on a regular basis.

The selected locations all contained FDM printers that are used on a regular basis. Additionally, the environments at all test locations were not modified in any way from their typical daily use and operations in order to best simulate the emission characteristics a printer operator or bystander might experience “in the wild” when in the same room as an operating FDM printer. Samples were taken to compare changes and trends during print time for total particle counts, fine particle concentrations, and VOC concentrations in the seven test environments.

Samples were taken to compare changes and trends during print time for total particle counts, fine particle concentrations, and VOC concentrations in the seven test environments. Fine particle size distributions were also analyzed and used to calculate fine particle mass concentrations to compare against National Ambient Air Quality Standards (NAAQS) exposure limits and VOC concentrations against National Institute for Occupational Safety and Health (NIOSH) exposure limits.Test Environments and Experimental Set UpTo test the impact of emissions of 3D printing “in the wild”, air sampling instruments were set up in seven different locations at the Penn State University – University Park campus where desktop 3D printers were in frequent use for class projects or research. The testing environments were not altered in order to best replicate the conditions of everyday use.

The testing environments were not altered in order to best replicate the conditions of everyday use. Table 1 summarizes the testing environment and material printed in each sampled location.Location 1 is a secluded room outside of a staff office designated for 3D printing forMechanical & Nuclear Engineering students. Students can schedule prints for class and project use. Locations 2 and 6 are printers in classrooms in the Engineering Design program where the University’s 3D Printing Club meets and prints various objects for club use and course projects. Location 3 is a secure room located in the University’s Pattee/Paterno Library, with a separate, sealed enclosure for all of the printers inside the room. Students upload 3D prints through the University’s library website and pick up their prints at the front desk in the library. Location 4 is a student dorm located on campus, where a student uses the printer for class and recreational use.

Location 4 is a student dorm located on campus, where a student uses the printer for class and recreational use. Location 5 is a small room located next to a workshop that is primarily used for wood and metal fabrication. Location 7 is a small graduate research office located nearby off-campus.2457Table 1.

The first is the TSI 8525 P-Trak Ultrafine Particle Counter which is capable of measuring fine particle counts ranging in size from 0.02 micrometer to 1 micrometer using a sample flow rate of 100 cm3/min and a total flow rate of 700 cm3/min. The P-Trak was lightweight and easy to position within the room. The P-Trak also had a flexible probe attachment that allowed for measurements to be taken in close proximity to the 3D printer and in confined regions larger than the P-Trak unit. The second instrument used to measure the coarser particle emissions was the TSI 9110 Aerotrak Portable Particle Counter with a size range of 0.1 to 10 micrometers and a flow rate of 28,000 cm3/min. The Aerotrack separates particles into 8 different bins based on particle diameters.

The Aerotrack separates particles into 8 different bins based on particle diameters. To measure VOCs, the RAE Systems PPBRAE Plus Photoionization Detector (PID) was used to count concentrations of gases with a range of 0-9999 parts per billion at a flow rate of 400 cm3/min.The P-Trak 8525 and the Aerotrak 9110 were first shipped to the manufacturer, TSI Inc,for calibration and qualification to ensure their specific ranges and sensitivities could be met. Both meters passed calibration by their manufacturer, TSI Inc., without issue. Following the manufacturer’s recommendations, isobutylene gas was used to calibrate the PPBRae PID, according to the manual’s specifications. Before any measurement was taken, the PID was zeroed using a VOC filter. A VOC filter ensures that the meter is reading zero VOCs when in a VOC free environment.

Before any measurement was taken, the PID was zeroed using a VOC filter. A VOC filter ensures that the meter is reading zero VOCs when in a VOC free environment. Likewise, the P-Trak meter was zeroed before every measurement using a high-efficiency particulate arrestance (HEPA) filter to verify the meter read zero particles in a particle free environment. In some test locations, printers were operating before our arrival, which complicated establishing control level VOC and PM measurements.During printing, the Ptrak Ultrafine Particle Counter was set next to the 3D printers withthe probe positioned close to the print bed. The instrument sampled air for the duration of the2458prints at five minute intervals, providing data on particle counts/cm3 at that instant. Meanwhile, the Aerotrak Particle Counter was placed within a foot of the 3D printer, and it sampled air continuously throughout the print.

Meanwhile, the Aerotrak Particle Counter was placed within a foot of the 3D printer, and it sampled air continuously throughout the print. The Aerotrak provided data on total particle count throughout the print and the total flow through the instrument. The PPBRae Photoionization Detector was handheld close to the printer for five minutes during every print to sample the VOC concentration. The PID took 5 readings, each with a 60 second sample, and calculated the minimum, maximum, and average concentrations of total volatile organic compounds in ppb.Calculating Fine Particle ConcentrationsParticle concentration data from the P-Trak Ultrafine Particle Counter was used tocalculate the number of fine particles (less than 2.5 micrometers in diameter) emitted in each environment.

While the P-Trak only provided particle count data for particles less than 1.0 micrometer in diameter, the Aerotrak particle counter indicated that the concentration of particles between 1 and 2.5 micrometers was insignificant in all test environments. With this in mind, the total concentration of particles measured using the P-Trak was used as the total fine particle count. Particle distribution data from the Aerotrak showed that more than half of the total particles counted during each sample fell between 0.1 and 0.15 micrometers in diameter. Table 2 contains particle distribution data from the Aerotrak for PLA-2 and Nylon-6. A plot of the data contained in Table 2 can be seen in Figure 1. Particle concentrations tend to be skewed towards a smaller diameter, moreso than a normal distribution would suggest because particle concentrations typically display a log-normal distribution behavior.

Because approximately half of the counted particles were greater than 0.15 micrometers and the remaining half were less than 0.15 micrometers, a Dp-50, or median value of 0.15 micrometers was used as the median diameter for fine particles.Table 2. Particle Collection Rate (# per minute) for PLA-2 and Nylon FilamentParticle Size (μm)Particles Emitted per MinuteBinPLA-2Nylon-60.105994825754600.152493813446960.20990581197390.251053331239690.3062504687620.50657857831.00288322275.003687292459Figure 1.

Assuming that the particle distribution of the Aerotrak was similar to the particle distribution of the P-Trak count, the same standard deviation was used for Equation 1.Using these values, a median particle diameter based on mass of 0.22 micrometers was calculated for fine particles. To calculate mass concentrations from fine and ultrafine particle concentrations, Equation 2 was used where c is equal to mass concentration and y is equal to particle concentrations.(cid:1855) (cid:3404) (cid:1877) (cid:3400)(cid:2869)(cid:2874)(cid:2025) (cid:3400) (cid:2024) (cid:3400) (cid:1830)(cid:3043),(cid:2873)(cid:2868)(cid:4666)(cid:1865)(cid:1853)(cid:1871)(cid:1871)(cid:4667)(cid:2871)Calculating Total Particle Counts and ConcentrationsThe Aerotrak particle counter counted the total number of particles per bin. To calculate the total number of particles, the sum was taken from these bin counts.

To calculate the total number of particles, the sum was taken from these bin counts. To calculate the particle concentrations in particles/cm3, the total particle counts for each bin were divided by the total centimeters of air that flowed through the meter during the sampling time. The concentration for each bin was also summed for a total. To compare particle counts across the seven different 3D2460(1)(2)printing environments, the total counts were divided by the sampling time (in minutes) to determine the total particle counts sampled per minute in each environment.Classifying Volatile Organic Chemical ConcentrationsAs mentioned in the previous section, isobutylene was used as the calibration gas for thePID for each sample taken. The PID has different sensitivities for different chemicals; thus, it gives different concentration readings for a certain chemical depending on which gas was used to calibrate it.

The PID has different sensitivities for different chemicals; thus, it gives different concentration readings for a certain chemical depending on which gas was used to calibrate it. To account for this, a correction factor was used for each chemical being detected to account for the PID’s sensitivity to that chemical relative to the PID’s sensitivity to isobutylene.

Each PID reading comprised of a minimum, median, and maximum VOC concentration in ppb. The average of the median and maximum values from the multiple PID readings was calculated for each environment sampled.The PID was only used to test total VOC count, not specific VOC composition. With that in mind, a composition of 60% lactide was assumed PLA-1, PLA-2, and PLA-3 (i.e., test environments 1, 2, and 3, which all used PLA). For ABS-4 and ABS-5, referencing VOC emissions data by MakerBot, a composition of 60% styrene and 40% unknown was assumed. For Nylon-6, referencing emissions data from Lulzbot manufactured nylon, a 90% caprolactam composition was assumed. For Polycarbonate-7, 80% of the emissions were assumed to be caprolactam [5].Experimental Results of 3D Printing “In the Wild”Fine Particle ConcentrationThe fine particle count and mass concentrations from the P-Trak ultrafine particle counterare summarized in Table 3.

Each testing environment shows an increase in average particle concentration during print time. The environments printing with PLA plastics had the lowest increase in average particle concentrations, with concentrations during printing increasing by 917, 157, and 465 particles/cm3 for Environments 1, 2, and 3, respectively. Environments printing with ABS and nylon plastic increased by 96,573, 13,850, and 35,931 particles/cm3 for Environments 4, 5, and 6, respectively. Environment 7, printing a polycarbonate blend, showed the highest increase in average particle concentrations with an overall increase of 170,597 particles/cm3. Maximum particle concentrations showed significantly higher increases over control concentrations. Environment 1 (PLA) showed a 2,094 particles/cm3 increase, Environment 4 (ABS) showed a 102,588 particles/cm3 increase, and Environment 7 (polycarbonate) showed a 299,927 particles/cm3 increase.2461(3)Table 3.

Fine Particle Count and Mass ConcentrationParticle concentration increases during printing with nylon filament were an order ofmagnitude higher than concentration increases for PLA filament, by a factor of 39.2. The particle concentration increase for Polycarbonate-7 was two orders of magnitudes higher than the concentration increase for PLA-1, as shown in Figure 2. Maximum particle concentration showed a significant increase over average particle concentration during printing for ABS, nylon, and polycarbonate filaments. For Environment 5 (ABS-5), maximum concentration was greater than average concentration by a factor of 2.33, Environment 6 (Nylon-6) showed a 1.86 factor increase, and Environment 7 (Polycarbonate-7) showed a 1.76 factor increase.

PLA usage increased concentrations by a factor of 1.60, 1.04, and 1.06 for Environments 1, 2, and 3, respectively.2.5) 3 ^ mr e p m a r g o r c i m( n o i t a r t n e c n o C s s a M2.01.51.00.5ControlAverageMaximum0.0PLA ‐ 1 Thermoplastic Filament ‐ Test EnvironmentPolycarbonate ‐ 7Figure 2. Fine Particle Mass Concentrations for PLA-1 and Polycarbonate-7 Test Environments2462For each test environment, fine particle concentration increased with time. Figures 3 and4 show fine particle concentration trends that appear to increase with time for Environment 1, which used PLA, and Environment 7, which used polycarbonate. PLA-1 shows a steady upward trend. Polycarbonate shows fluctuations (mostly due to interruptions during printing), while still trending upwards. Given the trends, if the prints were to have run longer, particle mass concentrations could be expected to increase at a steady rate.Figure 3.

Given the trends, if the prints were to have run longer, particle mass concentrations could be expected to increase at a steady rate.Figure 3. Fine Particle Concentration versus Print Time for PLA-1 Filament.Figure 4. Fine Particle Concentration versus Print Time for Polycarbonate-7 Filament.2463Fine Particle Mass Concentration and Size DistributionMass concentration is calculated directly from particle concentration; therefore, anincrease in mass concentration during printing is directly proportional to an increase in particle concentration during printing. A plot of fine particle mass concentration for all seven test environments can be found in Figure 5. Particle size distribution, which is correlated to mass distribution, was previously shown in Figure 1. Particle size distribution is skewed towards smaller average diameters. Approximately half of all particles measured in each test environment are less than 0.15 micrometers in diameter.

Particle size distribution is skewed towards smaller average diameters. Approximately half of all particles measured in each test environment are less than 0.15 micrometers in diameter. Table 2 showed that 53.3% of particles for Environment 2 (PLA-2) have diameters ranging from 0.10 to 0.15 micrometers. Moreover, 46.3% of particles in Environment 6 (Nylon-6) fall into the aforementioned diameter range. Particle diameter and particle count are inversely proportional to each other; when particle diameter decreases, particle count increases. Few particles in each test environment are greater than 0.5 micrometers in diameter.

Few particles in each test environment are greater than 0.5 micrometers in diameter. The percentage of total counted particles ranging from 0.5 to 5 micrometers for PLA-2 and Nylon-6 was 0.84% and 0.65%, respectively.2.5n o i t a r t n e c n o C s s a M e l c i t r a P e n i F) 3 ^ mr e p m a r g o r c i m(2.01.51.00.50.0PLA ‐ 1PLA ‐ 2PLA ‐ 3ABS ‐ 4ABS ‐ 5Nylon ‐ 6Polycarbonate ‐ 7Thermoplastic Filament ‐ Test EnvironmentFigure 5. Fine Particle Mass Concentrations for All Test LocationsTotal Particle CountWhen compared against control data, total particle count per minute does not change assignificantly as total fine particle count per minute. The relatively small change between control data and total particle count per minute for all seven test environments is shown in Figure 6. The largest total particle count per minute increase compared to control data occurs in Environment 7 (Polycarbonate), increasing by a factor of 1.1.

The largest total particle count per minute increase compared to control data occurs in Environment 7 (Polycarbonate), increasing by a factor of 1.1. This increase is two orders of magnitudes lower than the 437x increase found between control data for average fine particle count per minute for the same filament. Total particle count per minute in Environment 3 (PLA-3), when compared against control data, decreased by a factor of 0.99 during printing. The remaining total particle2464ControlAveragee t u n mir e p t n u o C e l c i t r a Pla t o Tcounts per minute increased by multiplying factors ranging from 1.0 and 1.1. When 3D printing, the largest total particle count per minute variation was between PLA-1 and Polycarbonate-7: total particle count per minute for PLA-1 was 53.6% higher than total particle count per minute for Polycarbonate-7.

Because the scale for total particle count per minute and average fine particle count per minute differ significantly, a comparison of the two would be insignificant.1,400,0001,200,0001,000,000800,000600,000400,000200,0000PLA ‐ 1PLA ‐ 2PLA ‐ 3ABS ‐ 4ABS ‐ 5Nylon ‐ 6Polycarbonate ‐ 7Thermoplastic Filament ‐ Test Environment Figure 6. Total Particle Count for All 7 Test EnvironmentsVolatile Organic Compound EmissionIn all 7 test environments, total VOC concentration increased during printing. Figure 7shows the trend of increased total VOC concentration during 3D printing. Environments printing with PLA and ABS were all towards the lower end, staying under 0.75 ppm. Nylon-6 reached a maximum of 2.96 ppm, and polycarbonate-7 reached a maximum of 2.57 ppm. Total VOC emissions for both nylon and polycarbonate were more than three times larger than the maximum VOC concentrations for both PLA and ABS.

Total VOC emissions for both nylon and polycarbonate were more than three times larger than the maximum VOC concentrations for both PLA and ABS. While total VOC emission is a good indicator of overall emission characteristics, a more telling indicator is the primary emitted VOC. Table 4 shows the control, average, and maximum VOC concentration for the primary VOC emitted in each test environment. The numerical values in Table 4 are not as important relative to each other as they are to the exposure limits for the respective primary VOC emitted in each test environment. Exposure limits for the primary VOCs emitted in each test environment are discussed in the next section.2465ControlDuring PrintFigure 7. Total VOC Concentration Levels in Each Test EnvironmentTable 4.

Total VOC Concentration Levels in Each Test EnvironmentTable 4. VOC Concentrations for Maximums VOCs Emitted at each Test Environment [5]DiscussionOne of the main areas of concern with respect to emissions from material extrusionprinters, particularly FDM systems, is fine particles. Fine particles are a criteria air pollutant (CAP) and are respirable [8]. While exact health effects of inhaling fine particle emissions from 3D printers are unknown, fine particle emissions from other sources have been shown to be carcinogenic and cause major health problems [12]. ABS, nylon and polycarbonate, all of which are commonly used thermoplastic filaments, are known to release the carcinogenic VOCs styrene and formaldehyde when melted [13]. The National Ambient Air Quality Standards (NAAQS) has set exposure limits for fine particle emissions at thirty-five µg/m3 during a 24-hr exposure period and twelve µg/m3 for consistent exposure over the course of one year [9].

Of the tested filaments, the most significant emitters of fine particles were polycarbonate, ABS, and nylon. While maximum concentrations for all three do not exceed twelve µg/m3, a longer print time could lead to a higher maximum concentration. Figure 8 contains a plot of particle mass2466concentration, in micrograms per cubic meter, for Polycarmonate-7 with respect to printing time. If polycarbonate filament were used to print over eight hours, which is not uncommon, particle mass concentrations could increase from two to above twelve micrograms per cubic meter.Figure 8. Particle Mass Concentration versus Print Time for Polycarbonate-7 FilamentAnother area of concern is VOC emission. When collecting emissions data, Environment6 was found to have hazardous VOC concentration levels from using nylon thermoplastic filament in an inadequately ventilated room. Environment 6 had an average VOC concentration of 1.1 ppb and a maximum VOC concentration 2.66 ppb.

Environment 6 had an average VOC concentration of 1.1 ppb and a maximum VOC concentration 2.66 ppb. The primary VOC emitted during FDM printing with nylon filament is caprolactam [5]. VOC concentration levels in Environment 6 are higher than the NIOSH short term (ST) and time weight average (TWA) exposure limit of 0.660 ppm and 0.220 ppm, respectively, for caprolactam. Inhalation of caprolactam can cause dizziness, headaches, nosebleeds, vomiting, liver and kidney damage, and confusion [10]. It has also be documented that caprolactam inhalation can cause skin and eye irritation, as well as central nervous system damage [11].Total VOC emissions for ABS filament tested in Environment 5 came to 0.336 ppm and0.605 ppm for average and maximum concentration levels, respectively.

The primary VOC emission for ABS, styrene, did not exceed the NIOSH styrene exposure limit of 100 ppm TWA or 600 ppm ST [5, 11] in Environment 5; however, some of the organic chemicals comprising the total VOC emission remain unknown. ABS is known to emit formaldehyde, and while we do not know the exact concentration of formaldehyde emitted from these printers, we cannot conclude that the amount is safe (NIOSH limits for formaldehyde are very low, at 0.016 ppm2467TWA and 0.100 ppm ST). Only 2.4% of the maximum VOC emissions would need to be composed of formaldehyde to exceed the TWA limit.What made this experiment unique was that testing was conducted in open, or natural, printing environments, what we call 3D printing “in the wild”. Similar experiments have been conducted at other institutions in more controlled environments using either a flux chamber or some other form of contained testing environment [1, 5, 10].

Similar experiments have been conducted at other institutions in more controlled environments using either a flux chamber or some other form of contained testing environment [1, 5, 10]. The goal of testing in natural environments was to gauge what emissions are like in everyday FDM printing environments, specifically at locations around a large college campus that offers many 3D printing options. Environment 5, a student dorm, and Environment 7, a graduate research office, are great examples of unregulated environments. Both locations are small rooms with no purpose built ventilation to accommodate FDM printers. Additionally, both locations used filaments well- known for emitting harmful VOCs and high levels of fine particles. An indicator that a plastic filament is emitting high concentrations of VOCs or fine particles is odor.

An indicator that a plastic filament is emitting high concentrations of VOCs or fine particles is odor. Nylon, ABS, and especially polycarbonate filament had strong, distinct odors, which can be used as indicators that a specific filament is emitting fine particles or VOCs at a high rate.Closing RemarksThis paper presents emission data and statistics for multiple thermoplastic filaments usedin natural environments or 3D printing “in the wild”. Air quality testing in natural 3D printing environments was done in an attempt to replicate what a printer operator or bystander might be exposed to in terms of PM and VOCs. PM and VOCs emitted during use of PLA filament were relatively low when compared to ABS, nylon, and polycarbonate filament. VOC emissions from nylon were significantly higher than the exposure limit set by NAAQS for caprolactam [9].

VOC emissions from nylon were significantly higher than the exposure limit set by NAAQS for caprolactam [9]. Fine particle concentrations for polycarbonate did not exceed NAAQS exposure thresholds during testing [9]; however, longer print jobs could push concentration levels above the NAAQS exposure threshold, which would result in a hazardous operating environment for the users and bystanders.When operating a printer with nylon or polycarbonate filament, certain safety proceduresshould be followed. The work area should be properly ventilated to help reduce VOC and PM concentration levels. Additionally, individuals in close proximity to the printer should wear a respirator to avoid inhalation of caprolactam. When operating a FDM printer in spaces without adequate ventilation, fans should be turned on and nearby windows should opened to allow for clean air to flow around the printer and circulate in the room.

Enclosed 3D printing spaces should make use of carbon filters to help reduce harmful emissions produced when printing. When educating students about 3D printing, particularly FDM systems, proper methods of ventilation and filtration should be discussed to make them aware of these safety issues.AcknowledgmentsThis work was made possible through financial support from Penn State’s Center forInnovative Materials Processing through Direct Digital Deposition (CIMP-3D). We would also like to thank Professor Paul Kremer from Penn State’s Architectural Engineering department for allowing us to use his laboratory equipment.2468Bibliography[1] N. Afshar-Mohajer, C.-Y. Wu, T. Ladum, and D. Rajon, “Characterization of particulatematters and total VOC emissions from a binder jetting 3D printer,” Building and Environment, 93, Part 2, 293-301, 2015.[2] Y. Kim, C. Yoon, S. Ham, J. Park, S. Kim, O. Kwon, and P.-J.

This offered a three-dimensional environment for complex and dynamic interactions that govern the cell’s behavior in vivo. Borate based bioactive (13-93B3) glass of different concentrations (10 to 50 weight %) was added to a mixture of PCL and organic solvent to make an extrudable paste. AD-MSCs suspended in Matrigel was extruded as droplets using a second syringe. Scaffolds measuring 10x10x1 mm3 in overall dimensions with a filament width of ~500 µm and pore sizes ranging from 100 to 200 µm were fabricated. Strut formability dependence on paste viscosity, scaffold integrity, and printing parameters for droplets of AD- MSCs suspended in Matrigel were investigated.1. IntroductionBone defects, resulting from trauma, cancer, arthritis, infection, or congenital skeletal abnormalities, contribute to major surgeries performed every year. Autologous bone graft is still considered as the golden standard for most applications but creates donor site morbidity [1, 2].

Autologous bone graft is still considered as the golden standard for most applications but creates donor site morbidity [1, 2]. Allografts avoid these issues but have limited availability, concerns over immunogenicity, and potential disease transmission [3]. Several types of materials, including biocompatible metals, bioceramics, and biopolymers are currently being investigated as candidates for synthetic grafts. Additive manufacturing of these materials has shown that complex and strong implants can be made to treat different regions of bone, including load-bearing bone [4-6]. However, engineered bone scaffolds have not been as successful as autologous grafts due to insufficient vascularization and reduced biomechanical function [7, 8].Borate based bioactive glasses are biocompatible, osteoconductive, and angiogenic.

In comparison to the more common silicate based bioactive glass, such as 45S5 or 13-93 glass, bioactive borate glass (13-93B3) has a higher reaction rate (5-10 times faster than silicate glasses); is more resorbable (60 to 70% wt. loss) in a few days to weeks; and is angiogenic, antimicrobial, and osteo stimulatory/conductive [9]. Controlling the size of the glass particles means the degradation rate of the glass can also be controlled to some extent. With the bioactive glass, there is also the potential use of dopants that could increase the positive biological effects, such as angiogenesis. In comparison, hydroxyapatite resorbs slowly and undergoes little conversion to bone-like material after implantation and provides no flexibility to tailor the material properties for the application. Biocompatible polymers such as polycaprolactone (PCL),1718provide strength and elasticity to scaffolds.

Biocompatible polymers such as polycaprolactone (PCL),1718provide strength and elasticity to scaffolds. PCL is one of the most widely used materials in 3D printing for biomedical applications because of its low cost and excellent rheological and viscoelastic properties [10]. Though PCL has a slow degradation rate (>2 years compared to few months for poly(lactic acid)/poly(glycolic acid) copolymers), a composite scaffold of 13-93B3 and PCL may provide the benefits of both 13-93B3 glass and PCL materials.Mesenchymal stem / progenitor cells (MSCs) have been used for cell therapy and in tissue engineering because of their ability to differentiate into multiple mesenchymal lineages in vitro, immune modulatory effects, and angiogenic capacity [11, 12]. MSCs have been isolated from several tissues, including the bone marrow (BM-MSCs), adipose tissue (AD-MSCs), and skin tissue [13-16].

MSCs have been isolated from several tissues, including the bone marrow (BM-MSCs), adipose tissue (AD-MSCs), and skin tissue [13-16]. The frequency of MSCs in adipose tissue is much higher than the more commonly studied source of bone marrow, yielding 100 to 500 times more cells per tissue volume [17, 18]. AD-MSCs have similar self-renewal abilities, common surface epitopes, growth kinetics, and cytokine expression profiles to BM-MSCs. With the addition of MSCs, the scaffold is expected to improve its biomechanical and biological properties in order to better repair the target tissue.Recent research has focused on creating living or cell-laden grafts for tissue engineering [19-21]. In such techniques, cells or cell aggregates are dispersed, typically in a hydrogel, and then deposited layer-by-layer and solidified either by thermal or chemical processes to form a scaffold.

In such techniques, cells or cell aggregates are dispersed, typically in a hydrogel, and then deposited layer-by-layer and solidified either by thermal or chemical processes to form a scaffold. Though such a scaffold holds promise in wound healing, drug delivery, and certain tissue engineering applications, bone repair requires a certain amount of mechanical integrity and controlled degradation of the scaffold which is difficult to accomplish. Traditionally, biopolymer scaffolds are fabricated using the fused deposition modeling process, where a polymer is melted and deposited thereby making it difficult to print cells alongside because of relatively high working temperatures [7].

In this study, we investigate the feasibility of printing a 3D scaffold using a two syringe system with a biopolymer/bioglass composite dissolved in an organic solvent as a scaffold material whilst simultaneously printing cells suspended in Matrigel, a gelatinous protein mixture representing basement membrane.2. Materials and Methods2.1 Preparation of PCL+bioactive glass compositePCL (Sigma-Aldrich, St. Louis, MO) was dissolved in chloroform (Sigma-Aldrich, St. Louis, MO) in a covered glass container with the help of a stirrer at ~50°C. The weight percentage of PCL was varied from 1:1 to 5:4 (grams of PCL to mL of chloroform) to determine the best possible ratio for printing. An appropriate ratio was established by visually inspecting the paste and through filament extrusion using a digital syringe dispenser.

An appropriate ratio was established by visually inspecting the paste and through filament extrusion using a digital syringe dispenser. Then, 13-93B3 borate bioactive glass (Mo-Sci Corporation, Rolla, MO) (nominal composition – 53% B2O3, 20% CaO, 12% K2O, 6% Na2O, 5% MgO, 4% P2O5 in weight percentage) of size less than ~20 µm was added to the PCL:chloroform mix in five different weight percentages in increments of 10, ranging from 10% to 50%. A magnetic stirrer was used to uniformly mix the composite paste and it was ensured that there was no glass particle precipitate before transferring the paste to a syringe. Each ratio was tested using a Loctite® digital syringe dispenser (Henkel North America, Rocky Hill, CT) at air pressures ranging from 10 to 50 psi with nozzle tips ranging from 110 to 500 µm.17192.2 Scaffold fabricationThe PCL/glass scaffold (10 mm x 10 mm) was printed with 0-90° orientation of the filaments in alternate layers, shown as schematic in Figure 1a.

Printing was performed with an assembled DIY 3D printer (Geeetech, Prusa I3 A Pro) which was modified to have two syringes connected to digital syringe dispensers and are computer controlled. The 3D printer set-up is shown in Figure 1b. The printing parameters such as filament spacing, layer height, printing speed, etc. were identified based on visual inspection and optical microscopic images after a single layer extrusion. Printing parameters such as needle tip size (260 µm) and printing speed (~ 8 mm/s) were uniform for all paste concentrations. Parameters such as air pressure and filament overlap are correspondingly modified for different pastes. Samples of 3D printed scaffolds were sputter coated with gold/palladium (Au/Pd) for 60 s before performing scanning electron microscopy (SEM). SEM (Hitachi S-4700 FESEM, Hitachi Co., Tokyo, Japan) images were taken to evaluate the surface morphology of scaffolds and internal structure of the filaments.Figure 1.

SEM (Hitachi S-4700 FESEM, Hitachi Co., Tokyo, Japan) images were taken to evaluate the surface morphology of scaffolds and internal structure of the filaments.Figure 1. (a) Schematic showing the fabrication of PCL+Bioglass filaments in 0-90 layer orientation with cells suspended in Matrigel (red dots), (b) 3D printer set-up with digital syringe dispensers2.3 Preparation and printing of AD-MSCsDifferent concentrations of Matrigel, diluted in phosphate buffered saline (PBS), ranging from 4 mg/mL to 10 mg/mL were printed using 10 psi and a 160 µm (30G) diameter nozzle tip to determine the ideal concentration for printing with cells. Frozen vials of approximately 1x106 AD-MSCs were obtained from three separate donors (LaCell, New Orleans, LA). Vials were unthawed, plated on 150 cm2 culture dishes (Nunc, Rochester, NY) in 25 mL complete culture media (CCM), and incubated at 37.5oC with 5% humidified CO2.

Vials were unthawed, plated on 150 cm2 culture dishes (Nunc, Rochester, NY) in 25 mL complete culture media (CCM), and incubated at 37.5oC with 5% humidified CO2. After 24 hours, the media was removed and adherent, viable cells were washed twice with PBS, harvested with 0.25% trypsin/ 1mM EDTA (Gibco), and replated at 100 cells/cm2 in CCM. Media was changed every 3 to 4 days. For all experiments, sub-confluent cells (≤70% confluent) between passages 2 and 6 were used.After determining that a concentration of 9 mg/mL Matrigel yielded the preferred droplet size, AD-MSCs at a concentration of 1x106 cells/mL were suspended in Matrigel. The printing parameters for Matrigel+cells were tested by designing a blocked experiment with air pressure as the blocked variable (10 and 20 psi). The other two variables considered were extrusion time1720(0.025 s and 0.035 s) and distance from the substrate (100 µm and 200 µm).

The other two variables considered were extrusion time1720(0.025 s and 0.035 s) and distance from the substrate (100 µm and 200 µm). A sample size of n = 3 was used for each droplet deposition with a specific set of parameters as shown in Table 1. When printing with Matrigel, the nozzle, cells, pipette tips, and Matrigel were all kept on ice until just before printing.Table 1. Experimental set-up to determine the Matrigel+cell droplet parameters (n = 3) Distance from substrate (µm) Dispensing time (s) Air pressure (psi)100 200 100 200 100 200 100 2000.025 0.025 0.035 0.035 0.025 0.025 0.035 0.03510 10 10 10 20 20 20 202.4 Degradation of PCL+bioactive glass and AD-MSCs distribution in MatrigelThe degradation of the PCL/glass composite material was studied using a thin sheet of the composite prepared by pouring the PCL/glass mixture on a polished glass plate and tape casted using a doctor blade set at a thickness of 600 µm.

The measured thickness of the dried film was 60±10 µm. Samples measuring 3 cm x 3 cm were cut from the sheet and kept in 50 mL of simulated body fluid (SBF) at 37°C. After the desired time intervals, the composite samples were removed from SBF and dried at 50°C for 12 h. The weight loss of each sample was measured and the SBF solution was analyzed by inductive coupled plasma spectroscopy (ICP) for boron, silicon, calcium, and phosphorus. To analyze the MSC distribution in Matrigel, 4',6- diamidino-2-phenylindole (DAPI) stain was used to image the Matrigel+cell droplets using fluorescent microscopy.2.5 Effect of Chloroform Evaporation on AD-MSC ViabilityThe effect of chloroform evaporation from the scaffold on the viability of the AD-MSCs was studied by printing three layers of the 30% glass composite on a two chamber microscope slide (Thermo Fischer Scientific, Rochester, NY) then printing a layer of AD-MSCs at a concentration of 10x106 cells per mL of Matrigel.

The composite was printed according to the parameters given in Table 2 and the AD-MSCs were printed at 10 psi, 0.035 s dispensing time, and 200 µm from the slide. Four droplets of cells suspended in Matrigel, about 0.4 mm in diameter, were printed on each horizontal filament of the scaffold.The Matrigel was allowed to polymerize at room temperature for 20 minutes then 1 mL of CCM was added. The slides were then incubated at 37.5oC with 5% humidified CO2. After 2 hours, the cells were stained according to the directions given with life technologies LIVE/DEAD Cell Imaging Kit (ref. R37601, Eugene, OR) and examined under a fluorescent scope. This was repeated again after 24 hours on a different set of scaffolds.17213. Results and Discussion3.1 Fabrication of PCL+bioactive glass compositeSingle layer tests: The weight percentage of PCL was varied from 1:1 to 5:4 (in grams of PCL to mL of chloroform) to determine the best ratio for fabricating the scaffold.

During the initial set of tests, different compositions of paste were extruded using a hand-held syringe and with the help of a digital dispenser by varying the nozzle tip and air pressure. An air pressure between 30 and 50 psi provided uniform extrusion of the PCL+chloroform mixture. The ideal ratio of PCL and chloroform was determined to be 5 g of PCL to 3 mL of chloroform, extruded at 30 psi using a 260 μm (25G) nozzle tip. A larger tip size (>260 µm) would result in thick filaments which are considered not beneficial for achieving faster scaffold degradation and smaller pore size distribution in the scaffold. Afterwards, 10% 13-93B3 glass by weight was added to the PCL+chloroform mixture and then extruded with the same set of parameters without difficulty. Figure 2a to 2c shows the filament extrusion tests performed on a microscopic glass slide with varying printing speeds.

Figure 2a to 2c shows the filament extrusion tests performed on a microscopic glass slide with varying printing speeds. A reduced filament width (from 1.8 mm to 0.8 mm) can be observed with increasing table speed, which was tested from 3 to 10 mm/s. The filament width is also dependent upon the homogeneity of the mixture. Figure 2d shows a successful and continuous single layer of a 15 cm x 15 cm scaffold with the filament spacing of 0.7 mm and a printing speed of 8 mm/s.1.0 mm0.8 mm1.8 mmFigure 2. Single layer tests with (PCL+10% Glass) and different printing speeds (a) 3 mm/s, (b) 5 mm/s, (c) 10 mm/s, and (d) 8 mm/s (which is used for rest of the experiments).Two layer tests: The filament height and spacing of the first layer would be crucial parameters to print successive layers. Filament height would determine the layer height and filament spacing would define how well the bridging occurs in successive layers.

Filament height would determine the layer height and filament spacing would define how well the bridging occurs in successive layers. Figure 3a shows the optical microscopic image of the cross-sections of the first layer filaments. The average height of the filament (shown between the two arrows) was ~75 µm. The height of the filaments for 40 wt% and 50 wt% compositions remained the same as long as the same nozzle tip was used. The roundness of the filament improved with a smaller tip but because of the nozzle clogging issues, all the experiments were carried out with a 260 µm tip. Therefore, a layer height of 0.08 mm was used to fabricate subsequent scaffolds. Another important factor in this study is the dwell time between consecutive layers as this allows the chloroform to evaporate thereby allowing the previous layer to become dry. A longer dwell time (>5 min) would warp the layer and a shorter dwell time (<1 min) is not sufficient for the layer to dry.

A longer dwell time (>5 min) would warp the layer and a shorter dwell time (<1 min) is not sufficient for the layer to dry. Figures 3b and 3c show the results of printing a second layer with no dwell time and with 1 min dwell time, respectively, with a 0.8 mm filament spacing. The difficulty in bridging the second layer without dwell time1722can be noticed, which was substantially improved with 1 mm dwell time. Figure 3d shows the bridging of second layer with 2.5 min of dwell time and 0.7 mm filament spacing.Figure 3.

Figure 3d shows the bridging of second layer with 2.5 min of dwell time and 0.7 mm filament spacing.Figure 3. (a) Cross-sectional view of the PCL/glass filament measuring ~75 µm in height, (b) second layer printing with zero dwell time, (c) second layer printing with 1 min dwell time, and (d) second layer printed on top of the layer shown in Figure 2d with 2.5 min dwell timeMultiple layer scaffolds: Based on the best printing parameters determined, scaffolds with multiple layers were fabricated with all five PCL/glass compositions (10 to 50 weight percentage of 13-93B3 glass). For compositions made with 40 wt% and 50 wt% glass, the amount of chloroform required to obtain a pourable characteristic for the paste was 4 mL (instead of 3 mL to 5 g of PCL). This is believed to be due to the increased viscosity of the paste with increase in glass content.

This is believed to be due to the increased viscosity of the paste with increase in glass content. With the exception of the paste made with 30 wt% of glass, which was extruded at an air pressure of 40 psi, remaining paste compositions were extruded at an air pressure of 30 psi. The final set of parameters used to fabricate scaffolds is given in Table 2. The filament width of 500±50 μm was measured for all compositions printed with a 260 µm nozzle tip while the average pore size depended on the filament spacing. A spacing of 0.6 mm provided pore sizes ranging from ~100 to ~200 μm while in scaffolds with a spacing of 0.7 mm, the pore size varied from ~200 to ~300 μm. Figure 4a shows an optical microscopic image of a scaffold fabricated with a filament spacing of 0.6 mm and with a smaller pore size distribution. As can be seen, pore sizes could be adjusted by modifying the filament spacing, to a certain extent, to suit a certain tissue engineering application of the fabricated scaffold.

As can be seen, pore sizes could be adjusted by modifying the filament spacing, to a certain extent, to suit a certain tissue engineering application of the fabricated scaffold. It is known that pore size is an important aspect of the scaffolds that could potentially effect the bone growth after implantation and it has been reported that pore sizes in the range of 100 to 300 µm are beneficial for bone growth [8]. The scaffolds fabricated by our process have pores in the same range. Figure 4b shows images of 10 mm x 10 mm scaffolds fabricated with 0.8 mm spacing.Table 2. Final printing parameters for each of five PCL/glass paste compositions13-93B3 Glass (Wt. %) 10 20 30 40 50Air Pressure (psi) 30 30 40 30 30PCL:Chloroform (g to mL) 5:3 5:3 5:3 5:4 5:4Filament Spacing (mm) 0.6 – 0.8 0.6 – 0.8 0.6 – 0.8 0.7 – 0.9 0.7 – 0.91723Figure 4.

%) 10 20 30 40 50Air Pressure (psi) 30 30 40 30 30PCL:Chloroform (g to mL) 5:3 5:3 5:3 5:4 5:4Filament Spacing (mm) 0.6 – 0.8 0.6 – 0.8 0.6 – 0.8 0.7 – 0.9 0.7 – 0.91723Figure 4. (a) Optical microscopic image showing the pore size distribution (~150 µm) in a scaffold fabricated with 0.6 mm filament spacing, (b) (L-R) pictures of representative scaffolds made with 10% to 50% glass content; the bottom image shows the warpage of scaffolds containing 10 to 30 wt.% glass with the red arrow indicating the space between the scaffold and slide. No warpage is seen in 40% and 50% scaffolds; scaffold with 50 wt% glass is the thickest fabricated (1 mm).It should be noted that the maximum thickness (or height) of the scaffolds depends on the degree of chloroform evaporation and the distance between layers. All of our experiments were carried out at room temperature (64°F) where the variation in relative humidity (58-60%) was not considered to be a major factor.

All of our experiments were carried out at room temperature (64°F) where the variation in relative humidity (58-60%) was not considered to be a major factor. Faster chloroform evaporation would produce warpage of the fabricated scaffold, especially with some dwell time between the layers. Non-uniform distribution of the PCL and glass is not believed to be one of the major factors of warpage as, upon examination of the filaments’ microstructure when printed with the same syringe at different time intervals, there was similar and uniform deposition of glass particles throughout the matrix. Therefore, the chloroform evaporation and the percentage of PCL in the composite is one of the crucial factors which determines the warpage. Increasing the glass content in the composite would indirectly decrease the chloroform content and thereby aids in faster evaporation and improves the filament rigidity.

Increasing the glass content in the composite would indirectly decrease the chloroform content and thereby aids in faster evaporation and improves the filament rigidity. In this study, it was observed that after reaching a thickness of about 0.64 mm (8 layers), the scaffolds being fabricated with 10% and 20% glass exhibited warping which led to difficulty in printing successive layers (see Figure 4b). The warpage in scaffolds made with 30% and 40% glass was less pronounced and thickness of 0.8 mm (10 layers) was obtained. The best results were achieved for 50% glass scaffolds as they were successfully printed to 1 mm thickness (12 layers) and could possibly have successive layers printed. Larger thicknesses were not attempted as the focus of this study is on the feasibility of printing PCL/glass scaffold along with MSCs. Though the mechanical properties of the scaffolds are not measured, it was observed that the ease of handling scaffolds improved with increasing glass content.

Though the mechanical properties of the scaffolds are not measured, it was observed that the ease of handling scaffolds improved with increasing glass content. The scaffold made with 50% glass had enough strength to be safely handled.Microstructure of PCL/glass composite scaffolds: Figure 5 shows the scanning electron micrographs of a couple of representative PCL/glass scaffolds made with 40 and 50 wt.% glass. Figures 5a to 5c show the surface morphology of the filament with increasing magnification (x30, x90, and x2000). It was interesting to observe that particles of bioglass are conspicuously absent from the surface of filaments. No pores on the filament surface were detected even when observed at a high magnification of x2000. Figures 5d to 5e show the filament fracture surface with increasing magnification (x180, x1000, and x2000). Glass particles dispersed in the PCL matrix can be seen in the interior.

Figures 5d to 5e show the filament fracture surface with increasing magnification (x180, x1000, and x2000). Glass particles dispersed in the PCL matrix can be seen in the interior. The dissolved PCL in chloroform encloses the glass particles1724and the surface tension effects between the steel nozzle tip and PCL during extrusion might cause the presence of only PCL on the surface.Figure 5. SEM images of the PCL/glass scaffolds. (a-c) Images with increased magnification from L-R showing a smooth surface morphology of the filament (40% glass scaffold), (d-f) Images with increased magnification from L-R (50% glass scaffold) (d) showing the fracture surface, (e-f) porous cross-sectional area of the filament with PCL matrix and glass particles3.2 Degradation of PCL+bioactive glass composite in Simulate Body FluidA “PCL+13-93B3 glass+chloroform” composite system has been studied in the recent past by producing thin sheets (60±10 µm) of PCL/glass composite [22].

The 3 cm x 3 cm sheet samples were soaked in SBF, dried overnight and weighed to measure their weight loss. The results indicate that 13-93B3 glass in the PCL/glass sheet had fully reacted to form hydroxyapatite (HA) in about 3 days. Figure 6 shows the weight loss percentages of the composite sheets made with different glass contents (20B – 20 wt.%, 40B – 40 wt.%, and 50B – 50 wt.% of 13-93B3 glass). In all three composites, reaction of the borate glass occurred rapidly during the first three to six days and then weight loss remains nearly constant. As expected, there was no loss of weight after 14 days for PCL. Also, weight loss increases with the increasing weight percentage of borate glass. The arrows on the right axis indicate the ideal weight loss for each composite which is weight loss of the glass if it completely reacted in SBF to form stoichiometric HA (Ca10(PO4)6(OH)2).

The arrows on the right axis indicate the ideal weight loss for each composite which is weight loss of the glass if it completely reacted in SBF to form stoichiometric HA (Ca10(PO4)6(OH)2). A similar degradation profile for 3D printed composite scaffolds with different glass compositions is expected.1725Figure 6. Weight loss for 50%, 40% and 20% glass compositions w.r.t weight loss of 100% PCL. The arrows on the right indicate the ideal weight loss for each composite [22].3.3 Dispensing AD-MSCs suspended in MatrigelExperiments were conducted to print droplets of Dulbecco's Modified Eagle Medium (DMEM) which contained suspended AD-MSCs. It was determined that a 110 μm (32G) nozzle tip extruded droplets less than 500 µm suitable for printing either on top or alongside the deposited PCL/glass filaments. However, the DMEM in the printed droplets would evaporate quickly making it difficult for further investigation.

However, the DMEM in the printed droplets would evaporate quickly making it difficult for further investigation. Therefore, the option of using Matrigel as the medium to suspend the MSCs was considered. The initial set of experiments included dispensing the Matrigel droplets without cells with the syringe dispensing system set-up to determine an appropriate concentration of Matrigel and droplet size. A concentration of 10 mg/mL Matrigel provided smaller drops (~100 μm), while 8 mg/mL Matrigel produced larger drops (~500 μm), and 4 mg/mL Matrigel produced even larger drops (1 mm). In each case, Matrigel provided a stable environment for the cells without drying (measured for up to 10 minutes). As the filament width of the scaffolds was measured between 400 to 500 µm, a Matrigel concentration of 9 mg/mL was selected to be appropriate for generating droplets which could be deposited on top of the filaments. Approximately 1x106 cells suspended in PBS were pipetted in Matrigel.

Approximately 1x106 cells suspended in PBS were pipetted in Matrigel. The AD-MSCs+Matrigel solution was then transferred just before printing to a 160 μm nozzle tip which was stored on ice during the entire non-printing time. It is assumed that the MSCs were uniformly distributed in the Matrigel before the start of the droplet deposition.An experiment was conducted to dispense Matrigel droplets with suspended MSCs by varying parameters including distance of the nozzle tip from glass slide, dispensing time of droplet, and air pressure. ImageJ software was utilized to quantify the number of cells in each fluorescent image. Figure 6 shows DAPI stained images of Matrigel droplets with 1x106 cells/mL printed at different parameters. Figures 7a-7d show the fluorescent images of the1726Matrigel droplets printed with 10 psi air pressure and Figures 7e-7h show the Matrigel droplet images printed with 20 psi air pressure.

Figures 7a-7d show the fluorescent images of the1726Matrigel droplets printed with 10 psi air pressure and Figures 7e-7h show the Matrigel droplet images printed with 20 psi air pressure. It can be clearly observed that droplets made at higher air pressure have a blue ring (cells are identified by blue dots in Figure 7) indicating that cells are at the boundary of the droplet because of the high pressure. This result is irrespective of the other two parameters. Droplets printed at the low air pressure (10 psi) have a smaller diameter and provide a more uniform distribution of cells. Such a distribution would be beneficial and the droplet size (<500 µm) would be appropriate to print on the PCL/glass filament. Among the four sets of droplets printed with 10 psi, it is observed that those printed with a distance offset of 200 µm had more cells (150 and 153) in comparison to those printed with an offset of 100 µm (105 and 148).

Further, amongst the droplets printed with 200 µm offset distance and 10 psi pressure, the droplets printed with a pulse time of 0.035 s was measured to have slightly higher cell count (153) in comparison to those printed at 0.025 s (150). These results allow us to determine the printing parameters for depositing Matrigel droplets, which are: (i) air pressure of 10 psi, (ii) distance from glass slide of 200 µm, and (iii) pulse duration of 0.035 s. Figure 6a shows a Matrigel droplet printed with the above set of parameters. Select cells are marked using arrows for better comprehension of the image in print. Investigating the cell survivability on the “PCL+13-93B3 glass+chloroform” filament is a crucial step toward our goal of establishing a novel and successful method of 3D printing of scaffolds with living cells for tissue engineering.1727Figure 7. DAPI stained fluorescent images of AD-MSC/Matrigel droplets, printed at (a-d) 10 psi and (e-h) 20 psi air pressure.

DAPI stained fluorescent images of AD-MSC/Matrigel droplets, printed at (a-d) 10 psi and (e-h) 20 psi air pressure. (a) A pulse time of 0.035 s and 200 µm distance from glass slide (b) 0.025 s and 200 µm (c) 0.035 s and 100 µm (d) 0.025 s and 100 µm. The cells are at the boundary of the droplet because of higher air pressure.3.4 Effect of Chloroform Evaporation on AD-MSC ViabilityViability of AD-MSC after printing was determined by a live/dead assay 2 and 24 hours after printing (Figure 8). At two hours, 96% were viable, demonstrating a minimal negative effect on the cells shortly after printing. At 24 hours, 65% were viable, indicating the potential for long term growth of the cells.ab100 µm100 µmFigure 8. Live/Dead images of AD-MSC/Matrigel droplets printed on 3 layers of the 30%glass PCL composite. Imaged after (a) 2 hours and (b) 24 hours. The yellow square indicates a pore in the scaffold.17284.

Imaged after (a) 2 hours and (b) 24 hours. The yellow square indicates a pore in the scaffold.17284. ConclusionThis study investigated the feasibility of fabricating a scaffold with polycaprolactone (PCL) and 13-93B3 bioactive borate glass composite utilizing a 3D printer without any heat input. This method would allow the process to incorporate cells during the printing of a scaffold unlike other processes where heating of the biopolymer is involved. Scaffolds were printed and near optimal printing parameters for each of the five different PCL/glass compositions were determined. Scaffolds fabricated with a 50:50 (in weight percentage) PCL/glass composite utilizing the parameters of 30 psi, 5 g of PCL to 4 mL of chloroform, 0.8 mm filament spacing were easy to handle with sufficient mechanical integrity.

Printing parameters for depositing cells suspended in Matrigel were determined and uniform distribution of cells in a ~400 µm droplet size was obtained for an air pressure of 10 psi and 0.035 s pulse. A live/dead assay performed 2 and 24 hours after printing cells on a 3 layer scaffold showed minimal negative effects from chloroform evaporation on the cells. The results of this study show the potential of the process to fabricate a scaffold with living cells embedded for tissue engineering applications.A continuation of this study would include increasing the height of the scaffold. To achieve this, the scaffold fabrication process will be modified by avoiding the continuous printing of the single layer and incorporating start-stop operations to deposit each filament in the layer. Such an operation would avoid the excess build-up of material at the scaffold edge which aids in building thicker scaffolds. The objective of this study is to simultaneously print the MSCs and PCL+glass.

The objective of this study is to simultaneously print the MSCs and PCL+glass. Therefore, the viability of printing MSCs based on the height of the scaffold will also be investigated. The degradation of the PCL+glass scaffolds is a work in progress and despite its known hydrophobicity, the initial results indicate that DI water/cell culture media is able to penetrate PCL and helps in glass dissolution. In this context, it is believed that the fabricated scaffolds tend to be more porous than the designed porosity of ~50%. Further experiments will be performed to investigate the overall scaffold porosity. In relation to cell viability, a live/dead assay will be performed for 24 hours, 1 week, and 2 weeks with the 50% glass composition.AcknowledgmentThe 13-93B3 glass used in this study was provided by MO-SCI Corporation, Rolla, MO. The authors thank Wenbin Li for his help in setting up the printer and Mariahe Long for her help with imaging.References1. J.C. Banwart, M.A.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed Paper3D Printing of Complex Wire Geometries for Tailored Resistance Response Tim Phillips*, Jared Allison*, Joseph Beaman*Department of Mechanical Engineering, University of Texas at Austin, TX 78712AbstractAdditive manufacturing (AM) is a rapidly growing field that enables production of complex geometries without tooling. AM has gained traction as a method of producing complex electronic circuits not possible using traditional techniques. The method explored in this manuscript involves post-build infiltration of conductive inks into complex channels to create resistive elements with tunable properties. A Polyjet printer is used to enable high-precision multi- material components with custom mechanical properties. Further, the conductive pathway geometry can be designed to achieve different resistive responses.

Further, the conductive pathway geometry can be designed to achieve different resistive responses. These properties allow for decoupling of the stress-strain response and resistance-strain response to produce custom strain gauges with engineered properties.IntroductionAdditive manufacturing (AM) is the process of building parts by joining material, typically layer-on-layer, directly from CAD data [1]. While the field originated as a means of producing functional prototypes, it has experienced remarkable growth in recent years and today is used in a wide variety of prototyping, tooling, and end-use applications. One growing area of interest is in AM of electronics, enabling complex circuitry not possible with traditional circuit manufacturing methods. The primary hurdle in AM of electronics is achieving a high-resolution conductive pathway within the components.

The primary hurdle in AM of electronics is achieving a high-resolution conductive pathway within the components. AM has shown tremendous potential to create complex geometries, but integrating conductive materials and circuit elements still poses a challenge. Currently, the most popular techniques of incorporating conductive materials are wire embedding, direct ink writing, aerosol jet printing, and coating with conductive materials.Wire embedding yields pathways with excellent electrical properties dictated by the embedded wire but are limited to planar or simple surfaces [2]. There is also the issue of connecting the wire to an embedded circuit element, which must be done during the embedding process, such as through laser welding [3] or by clamping between conductive plates also embedded in the part [4].Direct ink writing involves extruding continuous filament or droplets of an ink.

This can be a self-contained process where the ink makes up the entirety of the part [5] or incorporated into another process where a conductive ink is written on a non-conductive substrate [6–8]. Each ink writing step takes place on a planar or simple surface; however, low fidelity 3D pathways can be created by stacking multiple planar extrusions. The out-of-plane resistance is expected to differ from in-plane resistance due to a smaller contact area and potential incomplete bonding between layers. Aerosol jet printing is similar to direct ink-writing, using an aerosolized nanoparticle suspension jetted at the substrate rather than a liquid ink.

Aerosol jet printing is similar to direct ink-writing, using an aerosolized nanoparticle suspension jetted at the substrate rather than a liquid ink. It has many of the same advantages of1681direct ink writing, but is capable of even finer detail; however, this makes it slower at building 3D parts [9,10].Coating with conductive materials involves direct application of conductive material, such as through screen printing, or creating a preferential plating area during manufacture that is subsequently selectively coated. Coating includes the application of conductive material as well as electroless plating conductive material onto the 3D printed part. Electroless plating has been used to successfully coat copper onto fused filament fabricated (FFF) and stereolithography parts, but so far has only been shown to create simple surface-level conductive pathways [11,12].

Conductive coating has been accomplished by creating channels within a part during manufacture then flooding with a conductive material, such as liquid metal [13] or nanoparticle ink [14]. This paper will utilize the conductive coating process due to the limitations highlighted in other methods of creating conductive pathways.This paper utilizes the material jetting (MJ) additive manufacturing process, also referred to as Polyjet. This type of AM system builds parts by selectively depositing liquid material droplets then solidifying them using ultraviolet light before building the next layer upon it. One big strength of MJ systems is their ability to deposit multiple materials with discrete or graded interfaces to create multi-material parts.

One big strength of MJ systems is their ability to deposit multiple materials with discrete or graded interfaces to create multi-material parts. The design complexity afforded by this process is utilized in this paper to create strain gauges with tunable responses to tension and torsion part deformation.Materials and MethodsSpecimens were printed using a Stratasys J750 Digital Anatomy printer using a variety of materials, given in Table 1. Multi-material parts were built with a discrete interface between materials. Two support materials were used: a standard gel-like support and an experimental liquid support. The liquid support was used when building the channels, allowing complex geometries without the challenges of removing the standard gel-type support. After building, parts were manually cleaned of support material via high-pressure water jet.Specimen renderings are provided in Figure 1.

After building, parts were manually cleaned of support material via high-pressure water jet.Specimen renderings are provided in Figure 1. Each specimen consisted of 4 parts: (1) rigid end caps, represented as a blue material in the renderings. (2) flexible gauge section, represented as clear in the renderings, (3) stainless steel pins for resistance measurement attachment, shown as silver circles in the renderings, and (4) a conductive channel that spans the flexible gauge section and connects the steel pins, represented as copper in the renderings. A total of four unique channel designs were created plus an additional design that was a combination of two, creating two distinct channels within the same part.

A total of four unique channel designs were created plus an additional design that was a combination of two, creating two distinct channels within the same part. The designs were: (1) a single straight channel, (2) five straight channels, connected within the rigid end caps to form one continuous channel, (3) a single helical channel, (4) a 3D stepped channel that only moved in one cardinal direction at a time in a winding pattern, dubbed a “low-poly helix”, and (5) two distinct channels within the same gauge section, a single line and a helix winding around it. The gauge section, defined as the region with flexible material, was nominally 25.4 mm long with a 10.9 x 10.9 mm cross-section, disregarding the void area of the channels. Each channel had nominally a 2 mm diameter. The rigid part has the same cross-section as the flexible part at their interface, then tapers down to a tab for insertion into the tensile testing machine.

The rigid part has the same cross-section as the flexible part at their interface, then tapers down to a tab for insertion into the tensile testing machine. Path lengths for the conductive traces in each specimen are given in Table 2.1682Table 1: Stratasys J750 materialsMaterial Name Properties Flexible Agilus30 Rigid Vero SUP706B Support Model Cleanser* Support *Model Cleanser is typically only used as a cleanser, but is available as a build material through the GrabCAD Print Beta software packagePurposeBuilding the flexible portions of the strain gauge Building the ends of the strain gauge, outside the gauge section Gel-like support for general purpose building Liquid support for strain gauge channelsFigure 1: Strain gauge designs featuring (left to right): single line, multi-line, helical, low-poly helical, and dual helical-line channels.

The blue material is rigid Vero, clear material is flexible Agilus30, copper material is MG Chemicals 843WB conductive paint, and silver material is stainless steel.Table 2: Specimen conductive trace lengthsSample Single-line Multi-line Helix Low-poly helix Dual helix-lineGauge Path Length (mm)25.4 127.0 124.0 122.9 151.6/25.4Measurement Path Length (mm)38.1 221.5 136.7 135.6 206.2/35.61683A water-based conductive paint containing silver-coated copper flakes (MG Chemicals 843WB) was used as the conductive media. It was applied via syringe to push through the strain gauge channel followed by a bolus of air so as to coat the channel with paint while leaving it hollow. Stainless steel pins (PEM TPS) were inserted into the channels prior to coating to create attachment points for resistance measurements. Post-coating, the parts were placed in a vacuum for 4+ hours to accelerate drying.

Post-coating, the parts were placed in a vacuum for 4+ hours to accelerate drying. Two coats of paint were used, drying after each coat.Resistance was measured via an NI myDAQ digital multimeter and recorded via LabVIEW. Measurements were taken every 0.5 s with 0.1 Ohm resolution and a measurement range of 0-200 Ohm.An Instron 3345 was used to perform extension experiments. A 5 kN load cell was used to measure force and the Instron stage movement was recorded as the strain gauge extension. The test consisted of pulling the specimen 2.5 mm at a rate of 5 mm/min. A picture of the experimental setup is provided in Figure 2. While force was recorded during testing and measured nominally as ~8 N at 2.5 mm extension, stress is not reported. The non-constant cross-section of the gauge section due to the channels, as well as the discrete interface between rigid and flexible material create stress concentrations that make stress calculations non-trivial.

Additionally, the focus of this paper is on tailoring resistance response for a given deformation, so the stress-strain response is not explored in this manuscript.Figure 2: Tensile test resistance measurement setup for the dual-channel specimen.A custom rotation rig, shown in Figure 3, was used to measure the resistance response of the samples under torsion. The rotation rig was manually operated, so it was not possible to torque1684the samples in a smooth, continuous fashion. Instead, each sample was twisted along its central axis positive and negative 90⁰ in 15⁰ increments and held for 15 seconds while the resistance was recorded.Figure 3: Rotation resistance measurement setup.Results and Discussion Tension ResponseThe resistance response of the strain gauges under tension are given in Figure 4 and Figure 5.

As expected, each design exhibited an increase in resistance as the specimen was stretched due to lengthening of the resistive element; however, the absolute value and linearity of the response was dependent on the geometry. At zero extension, the resistance value should be proportional to the measurement path length, with the shortest specimen, single-line, having the lowest resistance and the longest specimen, multi-line, having the highest resistance. This trend was observed; however, the separation between the helical specimens and multi-line specimen was not as great as expected despite the multi-line specimen having a significantly longer measurement path length. One possible explanation for this is that the conductive coating thickness varied between specimens, causing variations in resistance that are independent of measurement path length.

One possible explanation for this is that the conductive coating thickness varied between specimens, causing variations in resistance that are independent of measurement path length. Another possibility is that the amount and radius of curvature in the channels has an influence on the resistance.Best fit lines for the strain responses are provided in Figure 5. The single-line and multi- line designs both exhibited an exponential increase in resistance as they were stretched. The exponential multiplier of the multi-line specimen, 6.40, is approximately 5 times greater than the1685multiplier for the single-line specimen, 1.29, the same as the difference in gauge path length. Intuitively, this makes sense as the change in resistance should be proportional to the amount of resistive element being stretched. The helix also exhibited an exponential increase in resistance, though with a much smaller exponential value.

The helix also exhibited an exponential increase in resistance, though with a much smaller exponential value. The low-poly helix exhibited a more linear response.Figure 4: Resistance response for the strain gauge geometriesThree of the dual-channel specimens were printed with the same nominal geometry and tested under the same strain conditions, presented in Figure 6. The general trend of each specimen’s longer channel, the helix, having a higher resistance than the shorter channel, the line, remained true. There was some variation, however, when comparing specimens to each other. Specimens 1 and 2 were similar, despite the linear channel in specimen 1 having a stronger response than that of specimen 2, as shown by the steeper slope of the dashed blue line compared with the dashed red line in Figure 6.

Specimen 3, however, exhibited a drastically lower response in the helical channel (green solid line), so much so that the resistance of the linear channel (green dashed line) surpassed the helical channel resistance when stretched beyond ~1.0 mm. This response is more inline with what was observed in Figure 4, where the linear channel had a stronger response than the helical channels and had a higher resistance once stretched beyond a certain point. It is unclear whether the variations between these specimens is caused by differences in the coating or manufacturing process and further exploration is needed to determine how to produce more uniform and repeatable specimens.1686Figure 5: Resistance response of the strain gauges as a function of extension, δ.

Experimental data is shown in blue and the best-fit line is overlaid in red dashes with the equation printed on the figure.1687Figure 6: Resistance response under strain for the dual channel helix-line specimensTorsion ResponseAfter completing the extension experiments, the same samples were used in the rotation test rig. Between successive tests, each specimen was left in a relaxed state for at least 5 minutes, by which point the no-load resistance had returned to the initial level. The torsion response for the single-channel specimens is presented in Figure 7. A positive angle represents rotating the part counter-clockwise, which is aligned with the helix rotation direction. The helix, low-poly helix, and multi-line specimens all exhibit an increase in resistance when rotated in either direction.

The helix, low-poly helix, and multi-line specimens all exhibit an increase in resistance when rotated in either direction. The multi-line specimen is symmetric within the gauge section, so its resistance response is not expected to be dependent on the direction of rotation, only the absolute value of it. This is true up to a rotation of 60⁰, but beyond that, counter-clockwise rotation elicits an increased resistance response.The single-line specimen exhibits little-to-no resistance change regardless of rotation angle. This lack of response is expected since the resistive channel runs through the center of rotation for the torsion test rig, making it so there is no change in path length during rotation.The dual-channel specimens were tested in the same fashion, their results are presented in Figure 8.

The trends found in the single-channel specimens are also found in the dual-channel specimens; the longer helical channel has a higher resistance than the shorter linear channel and only the helical channel exhibits a resistance change when twisted. The resistance values vary significantly between specimens, further confirming that the coating was inconsistent across specimens. The helix in specimen 1 exhibited a significantly stronger response than in the other specimens when rotated counter-clockwise, so much so that rotating beyond 60⁰ resulted in a reading beyond the range of the myDAQ. The test was aborted after 60⁰ as helical data was no longer being recorded, but there was no indication that this affected the linear channel, which continued to show nearly zero change in resistance.

This test was repeated and the results were1688the same, where rotation beyond 60⁰ resulted in the helical channel resistance going beyond the measurement range.When released from tension or torsion and allowed to return to their natural state, the resistance of all specimens also returned to the original, un-flexed value. The resistance took a few minutes to return to the original value, likely attributable to the lossy nature of the Agilus30 flexible material used in the gauge section, which also took time to relax back to its natural state.Figure 7: Torsional strain response for single-channel specimens. A positive angle represents a counter-clockwise rotation, along the helical channel direction.Figure 8: Torsional strain response for the dual-channel Helix-Line specimens.

A positive angle represents a counter-clockwise rotation, along the helical channel direction.Figure 8: Torsional strain response for the dual-channel Helix-Line specimens. A negative angle represents a counter-clockwise rotation, along the helical channel direction1689ConclusionThis manuscript presents a novel low-cost process of creating multi-material 3D printed parts with conductive channels via post-build coating with conductive paint. The resistance of the channels was measured as the specimens underwent tension and torsion testing and showed that the resistance value is related to the channel design and type of part deformation. Based on these findings, it is believed that channels can be designed to achieve specific resistance responses, such as the resistance in a straight channel being unchanged by torsion around the channel axis.

This opens the door to multi-element strain gauges that can detect not only the amount, but also the type of deformation the part undergoes. Additionally, it is shown that the conductive coating is capable of bonding to both the 3D printed material as well as the inserted pins, indicating that other circuit elements can be inserted into the parts and connected to the circuit. Potential applications of this include part health monitoring, strain gauges with designed resistance response, and 3D circuitry with embedded circuit elements.There are a number of next steps identified that would further improve the utility of the process described in this paper. Developing a coating process that is consistent across specimens would improve the useability by making the resistance values predictable and controllable. Testing different materials and graded interfaces would further show the tuneability of the resistance response.

Testing different materials and graded interfaces would further show the tuneability of the resistance response. Lastly, embedding circuit elements within the polymer substrate mid-build and connecting them with the conductive pathways would open the door for functional electronics beyond strain gauges.Works Cited [1] 2017, ISO/ASTM52900-15 Standard Terminology for Additive Manufacturing - GeneralPrinciples - Terminology, ASTM International, West Conshohocken, PA.[2] Billah, K. M. M., Coronel, J. L., Halbig, M. C., Wicker, R. B., and Espalin, D., 2019, “Electrical and Thermal Characterization of 3D Printed Thermoplastic Parts With Embedded Wires for High Current-Carrying Applications,” IEEE Access, 7, pp. 18799–18810.[3] Espalin, D., Muse, D., MacDonald, E., and Wicker, R., 2014, “3D Printing Multifunctionality: Structures with Electronics,” The International Journal of Advanced Manufacturing Technology, (72), pp.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed Paper3D PRINTING OF COMPLIANT PASSIVELY ACTUATED 4D STRUCTURESDhileep Kumar Jayashankar1, Sachin Sean Gupta1, Loo Yi Ning Stella2, Kenneth Tracy31Research Assistant, 2Student, 3Assistant Professor Digital Manufacturing and Design Centre, Architecture and Sustainable Design Pillar Singapore University of Technology and Design, SingaporeAbstractAdditive manufacturing has begun to revolutionize the production of various physical technologies that depend on bespoke geometry and tailored material properties for function. This includes the design of compliant mechanisms, which rely on an integral coupling between geometric and material parameters to attain the elastic flexibility necessary to accommodate programmed deformation.

While kinetic structures with compliant parts are typically activated by the application of a mechanical force, alternative means of achieving motion are available, such as the use of smart, 4D, or stimuli-responsive materials which react to environmental conditions. In this research, a combination of compliant mechanisms and water-responsive chitosan biopolymers was explored to create flexible, programmable passive actuators, enabled by 3D printing. A set of compliant joints were modeled, simulated, fabricated, and tested to determine the optimal design for use in the actuator. The actuator was then iteratively tested with wetting and drying of chitosan films to invoke a specific shape change, which was analyzed for accuracy, speed, and consistency.

The actuator was then iteratively tested with wetting and drying of chitosan films to invoke a specific shape change, which was analyzed for accuracy, speed, and consistency. The study concluded with a discussion of the implications of synthesizing compliant mechanisms, chitosan biopolymer, and additive manufacturing for next-generation adaptive structures.Keywords: Compliant Mechanism, Passive Actuation, Additive Manufacturing, Chitosan BiopolymerIntroductionCompliant mechanisms (CM) leverage the elasticity of their flexible members to achieve motion [1]. This contrasts from traditional mechanisms, which are composed of assemblies of discrete, rigid parts connected at articulating joints. Dependence on material properties rather than mechanical components means that CM offer advantages such as no hysteresis, compactness, ease of fabrication, inherent simplicity, low weight, high reliability, and frictionless and wear-free motion [2].

These benefits indicate the CM can be implemented where typical maintenance for traditional joints becomes problematic. This appears in fields such as biomedical implants, soft robotics, building structures, space research, and engineering at micro- and nanometer scales [3][4]. While offering such appealing advantages, CM nonetheless suffer from design and fabrication issues: conventional manufacturing methods become infeasible for producing a CM with the bespoke geometry that may be required for a specific application. Advanced mechanism technologies are required for realizing novel compliant structures with specially programmed deformation.With its inherent freedom of geometric complexity and ever-expanding repertoire of materials to suit high- performance applications, additive manufacturing (AM) is poised to revolutionize that production of CM.

Products created through AM have rapidly transitioned in the last decade, from their humble beginnings as delicate scale models to novel, direct-use objects and structures with enhanced capabilities [5]. AM is a powerful tool for fabricating next-generation products whose performance requirements depend on both geometric and material parameters, which constitute an inseparable relationship in the case of compliant mechanisms. Mirroring the monolithic nature of compliant mechanisms, AM is reputed for its ability to fabricate objects in one continuous step. Furthermore, AM offers a heightened degree of customizability compared to other manufacturing processes, which in turn enables an expanded design space for CM. CM-AM has only just begun to be seriously investigated in variety of fields, spurred by recent advances in computer-aided design (CAD) and simulation tools. In the realm of soft robotics, Mutlu et al.

In the realm of soft robotics, Mutlu et al. have printed compliant prosthetic fingers using fused deposition modeling (FDM) of thermoplastic elastomers, following systematic finite element analysis (FEA) of different flexure types [6]. The team demonstrated that a single homogenous elastic material, when designed appropriately, could achieve the desired deflection behavior for the bending finger. Gaynor et al. opted for richer material exploration,948implementing multi-material inkjet printing to create a topologically optimized compliant force inverter [1]. This work proved the ability of AM to physically realize nonstandard CM design solutions, and the utilization of both rigid and soft photopolymers was key to improving the efficiency of the inverter. AM-CM has not only been restricted to polymers.

AM-CM has not only been restricted to polymers. NASA, the American aeronautics and space agency, has even experimented with titanium metal 3D printing using an electron beam melting (EBM) process to create lightweight, expandable parts which do not require lubrication; the authors emphasized that these properties are essential for space exploration [7].While the research developments mentioned above demonstrate the diversity of AM processes and materials that have been implemented in CM production, the common feature was the use of mechanical force to instigate the motion of the compliant structure. However, alternative sources of energy exist for inducing the compliant motion, which can be useful when mechanical inputs are neither desirable nor possible.

However, alternative sources of energy exist for inducing the compliant motion, which can be useful when mechanical inputs are neither desirable nor possible. Temperature- responsive, photosensitive, and hygroscopic materials have the ability to respond to environmental stimuli such as heat, light, and water, respectively, undergoing a change in physical state (such as shape or mechanical properties) that can be manipulated to achieve target motion. This opens the possibility to create smart, “4D” objects and structures that can transform independently when placed in a certain environment, with change over time representing the fourth dimension [8]. The term 4D printing refers to the direct printing of smart materials, whereas this work pursues an alternative strategy of printing the compliant parts that are later assembled with a smart material (chitosan).

Regardless of the approach, such structures are considered passively actuated due to a lack of active input of electrical or mechanical energy, assuming that the environmental stimulus is naturally present during use. This is especially poignant in the use of water as source of movement for passively actuated objects, given the ubiquity of ambient humidity, rain, or bodies of water. Promising research has already been conducted in the pairing of water-responsive materials with additively manufactured CM to create passive 4D objects. Foresterio et al. have demonstrated the use of multilayer wood veneer struts as cellulose-based actuators, joined with FDM-printed compliant polyurethane nodes to construct shape-shifting space frames that responded to variations in humidity [9]. In this case, the compliant nodes acted as bifold hinges that afforded degrees of flexibility to the struts, allowing them to rotate in order to morph into desired global curvatures. Tibbits et al.

In this case, the compliant nodes acted as bifold hinges that afforded degrees of flexibility to the struts, allowing them to rotate in order to morph into desired global curvatures. Tibbits et al. took a different route, directly printing hygroscopic polymers and nonactive polymers into compliant configurations, resulting in self-evolving structures when immersed in water [10]. These compliant structures took the form of printed assemblies of modules, links, and discs that could stretch or fold by leveraging differences in swelling capabilities to achieve targeted transformation.

While the results of both works showed the great potential of merging hygroscopic actuation with CM, limitations in general applicability resulting from slow response time (for the former) and load-bearing ability of the hygroscopic material in question (for both) were observed.However, alternative materials and compliant structure designs exist that can improve upon the current limitations in strength and speed of hygroscopically actuated compliant structures. So far to the authors’ knowledge, the combination of CM with chitosan biopolymer is yet to be explored. Chitosan is a derivative of chitin, the world’s second most abundant polymer after cellulose [11]. Chitosan biopolymer has been used in industrial applications ranging from cosmetics, agriculture, textiles, food processing, and tissue engineering.

Chitosan biopolymer has been used in industrial applications ranging from cosmetics, agriculture, textiles, food processing, and tissue engineering. Its extensive applicability stems from an exceptional combination of hygroscopicity, processability, and biodegradability, in addition to its ability to form composites with sound tensile strength [12]. These mechanical properties decrease appreciably when chitosan is exposed to water, causing amplified deformation to the structure to which chitosan is bonded [13]. This transition happens on the order of minutes, enabling rapid changes that can be accommodated by well-designed CM. Considering this potential technological synergy, this work explored the combination of CM and water-responsive chitosan biopolymers to create flexible, programmable passive actuators, enabled by 3D printing.

These actuators were afforded flexibility through the use of compliant joints that were designed to achieve a specific global curvature change of the actuator. In this study, different compliant joint designs were designed, modeled, simulated, printed, and tested. A single design was selected that was applied to the nodes of the actuator, which took the form of a truss-inspired cantilever fitted with hygroscopic chitosan films. This actuator was printed using fused deposition modeling (FDM) technology and tested for its motile capabilities in the pursuit of the target curvatures, including the repeatability of the transformation. These results were critiqued, as well as the general effectiveness of the synthesis of CM, chitosan biopolymer, and AM.949Design MethodologyTheory of compliant passive actuatorThe theory behind the motion of this passively actuated structure lies in strategic positioning of the chitosan films within the structural assembly.

Chitosan, a deacetylated form of chitin extracted from arthropods, is highly hygroscopic in nature and can be manufactured into different forms for engineering and medical applications [11]. In this work, chitosan is reinforced with cotton fiber gauze and formed into thin films. The production of chitosan films involves preparation of chitosan solution, filtering the solution with 50 μm accuracy to remove impurities, degassing the solution to remove air bubbles, and then casting the solution into molds. The films are then dried, neutralized, washed, and dried a second time. Based on measurements from previous work [12], these chitosan films have the ability to extend 12% in length with the absorption of water, thereby serving as a passive actuation mechanism.

Based on measurements from previous work [12], these chitosan films have the ability to extend 12% in length with the absorption of water, thereby serving as a passive actuation mechanism. Furthermore, chitosan films exhibit decreases in their mechanical properties when exposed to water [13]; thus, by having the chitosan films serve as critical tensile members, any loss in properties of the chitosan results in amplified deformation of the actuator. The cantilevering actuator purposefully fails when the chitosan becomes weak, eventually bringing forth a “wet state” configuration of the systems once the chitosan has fully elongated. As the chitosan films dry, they undergo significant internal forces that are sufficient to pull the actuator back upwards toward a “dry state” conformation once all water has evaporated. This operation can proceed in a cyclic manner and has been validated in previous studies using traditional revolute joints using a pin and ball bearing, as seen in Fig. 1.

This operation can proceed in a cyclic manner and has been validated in previous studies using traditional revolute joints using a pin and ball bearing, as seen in Fig. 1. This paper represents progress in attaining specific shape changes that can be programmed by the carefully designed deformation of compliant joints.I■ ■Fig. 1 Previous work of chitosan based passive actuator with revolute jointsSelection of Compliant Flexure DesignIn modern manufacturing, advanced mechanism technologies enable design engineers to use the concept of miniaturizing mechanical structures to achieve desired motion through new materials and with their possibilities by creating compliant structures. In this design methodology, a set of CM were explored using finite element methods (FEM) with FDM additive manufacturing technology to replace the conventional revolute joints with flexural hinges for actuator designs.

In this early stage of development, as opposed to following the pseudo rigid body approach or the topology optimization method to design the compliant flexure, intuitive physical prototyping was used to create a proof-of-concept. To achieve a planar motion, a primitive type single-axis flexure was designed as a three-node beam element with three degrees of freedom (DOF) per node in different geometrical configurations [2]. Various CM designs (henceforth abbreviated as CMD) were explored in terms of achievable range of motion based on different geometries and identical thickness to find the allowable stress and maximum deflection that were suitable for the structur(cid:68)(cid:79)(cid:3)movement provided by the chitosan film expansion. If950the flexure did not possess sufficient compliance, the movement of the actuator would be impeded.

If950the flexure did not possess sufficient compliance, the movement of the actuator would be impeded. This desired deflection was calculated as the vertical deflection of the most critical flexure between the neutral (partly wet and partly dry) and fully wet state actuator configurations, as shown in Fig. 2. Through measurement of the digital models, it was found that the joint deflection needed to reach 4.15 mm under the self-weight loading of the actuator to enable it to transition from initial state to wet state. The total truss weight in the dry state was estimated to be 20 gm. However, the standard weight of 50 gm which was nearest to the actual loading conditions was chosen to test the joint deflection. Therefore, a minimum threshold deflection under 50 gm was estimated, informed by FEM simulations of the joint deformation. This minimum deflection was coupled with a factor of safety of 1.5 to account for any underestimate in the simulation, resulting in a value of 9.34 mm.

This minimum deflection was coupled with a factor of safety of 1.5 to account for any underestimate in the simulation, resulting in a value of 9.34 mm. Thus, four different flexure designs were digitally and physically tested with 50 gm of weight hanging from their free end to reflect their loading behavior in the actuator. Their vertical deflections were measured and compared to 9.34 mm.- - - Wet state----- - ------joint deflection required! 4.15mm i9.34mmFEM + facrorofsaferyiunder actuator self-weightunder 50g test weightFig. 2 The flexure must be compliant enough to deflect 9.34 mm under the load from a 50gm test weight in order to achieve the targeted shape changeFour compliant models were designed based on attaining large angular deflection and with a low axis of rotation. The integral component of the CM in the four designs were flexures which allowed controlled bending that counted towards motion.

The integral component of the CM in the four designs were flexures which allowed controlled bending that counted towards motion. In traditional compliant designs, flexibility is attained along the axis by gradually thinning out material with specific geometry [2]. By assessing several primitives, it was observed at the initial scale of testing that the thinning out of material at one specific area introduces critical level of stress concentrations that eventually increases the overall stress, potentially causing the system to actuate inconsistently or even fail over a period of time after multiple cycles. To counteract this, only an effective design with minimal stress and larger angular deflection of flexure could overcome these challenges to achieve the desired motion in dynamic conditions. A series of flexure designs were conceived considering spring-based mechanisms to allow greater freedom of movement.

A series of flexure designs were conceived considering spring-based mechanisms to allow greater freedom of movement. The springy-derivative design brought two benefits: the load in each flexure could be spread out equally across a series of turns, instead of being constrained to one single area that weakens the structure; and secondly, it could be used to control the structure by defining pre-induced bending in the system. Fig. 3 shows the springy-derivative flexural designs.hilIUlIUlsi hIUUUlN ~ ~(a)(b)(c)(d)Fig. 3 Different springy-derivative flexural designs (a) CMD 1 (b) CMD 2 (c) CMD 3 (d) CMD 4951In addition to degree of motion, strength of the CM must be considered. Strength can be obtained by a load test to determine the force or the stress-deflection involved in the flexure. A static non-linear structural FEM analysis was applied to two-dimensional compliant flexure designs.

A static non-linear structural FEM analysis was applied to two-dimensional compliant flexure designs. The four compliant models, two with identical geometry but with different end conditions, and two alternatives, were analyzed by applying a point load of 0.5 N at the right extreme end, considering flexure as a cantilever beam. The material properties of the compliant flexure were an important characteristic to predict the fatigue failure through FEA methods. The properties of Stratasys’ ASA material in its XZ direction with a Young’s modulus and Poisson’s ratio of 2010 MPa and 0.394, respectively were utilized. Fig. 4 shows the maximum stress and maximum deflection of the four compliant mechanism designs. From FEM analysis, CMD 1 and CMD 2 did not possess significant levels of difference in stress with 47.843 MPa and 47.185 MPa, and similarly with deflections of 22.583 mm and 22.064 mm, respectively.

From FEM analysis, CMD 1 and CMD 2 did not possess significant levels of difference in stress with 47.843 MPa and 47.185 MPa, and similarly with deflections of 22.583 mm and 22.064 mm, respectively. CMD 3 and CMD 4 had a marginable difference in stress with 44.172 MPa and 46.189 MPa and deflections of 21.663 mm and 20.147 mm. The results were plotted between maximum stress and maximum deflection, as shown in Fig. 5, in order to find the suitable design which had low stress with high deflection. Out of the four designs, CMD 3 best satisfied the condition from an FEM simulation point of view. These simulation results were complemented with physical load tests of printed CMD.(g) CM03(h) CM04Fig.

Fig. 6 shows the results of these tests. CM Design 1 exhibited the highest deflection (15.1 mm) and CM Design 3 demonstrated the lowest (13.5mm); however, all the joint designs successfully surpassed the minimum compliance needed for the first joint. Thus, to finalize the flexure design, the stress results from the FEM simulation were the primary data used to narrow the selection. As stated above, CM Design 3 was predicted to undergo the least stress between all the designs, and so, CM Design 3 was selected for the actuator. The differences in the deformation results between the FEM simulations and the physical experiments can potentially be attributed to nuances of the AM process. In contrast to conventional manufacturing, AM-produced parts possess different properties in each direction that vary based on the direction of printing.

In contrast to conventional manufacturing, AM-produced parts possess different properties in each direction that vary based on the direction of printing. In this case, the CM were printed in the Z direction, which is perpendicular to the applied load, and so in the FEM simulation, only the tensile strength and Poisson’s ratio for the Z direction were included. This might have caused differences from the physical behavior. Additionally, because it is supported at only one end in the physical experiment, the CM has greater degrees of freedom to rotate or twist in other axes, which might have decreased the measured vertical displacement. In future studies, flexural strength and a tailored bending theory should be included in FEM to approximate the solution.Design of compliant structuresThe passive actuator assembly was designed as a spatial structure, inspired by the lightness of a space truss.

The assembly consists of structural members, CM Design 3 at the nodes, and chitosan films that act as critical tension elements. Through expansion and contraction of the chitosan films mediated by the compliant joints, a desired physical transformation of the structure could be achieved. The cotton-fiber reinforced chitosan film was responsible for the passive actuation due to their responsivity to moisture. A detailed research study about the actuation time for chitosan film was conducted by the authors and recently published [12]. One notable piece of information gained from the work was the use of cotton-fiber reinforcement for improved strength of chitosan film. Additionally, the orientation of the reinforcement plays a vital role by accommodating the films’ hygroscopic expansion. As shown in Fig.

Additionally, the orientation of the reinforcement plays a vital role by accommodating the films’ hygroscopic expansion. As shown in Fig. 7, 0°/90°-oriented films have restricted expansion due to horizontally- and vertically-aligned fibers, whereas 45°-oriented films possess greater expansion due to crisscross lay-up of fiber along their direction of applied force. From the authors’ previous work, it was reported that chitosan films with 45° oriented cotton-fiber reinforcement can expand up to 12.8 % linearly along the direction of applied force in a completely wet state (i.e. saturation state) due to hygroscopic swelling. Furthermore, the films yielded an953internal drying force of 19 N during drying in the reverse direction, which is the property enabling lifting of the actuator into its dried conformation.+-+--+-+--+-+---+--+-+---+---< ~ 1(a)(b)Fig.

7 Orientation of cotton-fiber reinforcement (a) 0°/90° with respect to the applied force (b) 45° with respect to applied force(a)(c)(b)954Production of AM partsThe actuator was printed on a Stratasys Fortus 450mc machine, which operates through fused deposition modeling (FDM) printing technology. FDM, one of the most widespread of AM processes, was selected for its ability to create parts of satisfactory strength and durability at low cost. The model of the passive actuator was converted into an .stl file and then imported into Stratasys’ Insight software for slicing as shown in Fig. 9 (a). A default layer height of 0.178mm using the T12 nozzle tip was selected, with choice of sparse infill for both the part interior and support material. The material selected for the actuator was ASA, Stratasys’ propriety model material that is similar to acrylonitrile butadiene styrene (ABS) in properties.

The material selected for the actuator was ASA, Stratasys’ propriety model material that is similar to acrylonitrile butadiene styrene (ABS) in properties. SR30 was selected as the corresponding water-soluble support material, enabling attainment of the future part with minimal post- processing. Upon finalization of the print settings, machine-readable files with toolpath instructions were automatically generated using the Insight software and sent to the printer. The actuator was printed as a single part within 4.5 hours and then placed in a solution bath to remove the support material as seen in Fig. 9 (b). In parallel, enclosure strips for securing the film on the actuator were laser-cut from thin acrylic sheets.____ 20 mmPrinted Compliant Design 3Fig.

9 (b). In parallel, enclosure strips for securing the film on the actuator were laser-cut from thin acrylic sheets.____ 20 mmPrinted Compliant Design 3Fig. 9 3D printing of the actuator (a) Sliced model of the actuator before printing (b) Print results of CM Design 3Assembly and Test SetupUpon completion of 3D printing and laser-cutting of chitosan films, the various components of the actuator were assembled. The films were sequentially laid upon the actuator, clamped with the acrylic strips, and secured with metal fasteners. The fully realized actuator first assumed a neutral flat conformation as opposed to a curled dried one due to exposure of the film’s due ambient humidity. The actuator was mounted to an acrylic mounting surface inside a test setup that was created to interchangeably administer water or dry air to the structure.

The actuator was mounted to an acrylic mounting surface inside a test setup that was created to interchangeably administer water or dry air to the structure. To test the performance of the compliant actuator, water and air were applied to the films for 20 minutes and 45 minutes, respectively, in a cyclic manner. The motion of the actuator was recorded using videography techniques. Three quantitative parameters, amount, speed, and repeatability of the deformation, were measured by analyzing and superimposing sequence frames from the videos. The physical results were then compared to the designed curvatures to verify the latter’s accuracy.Results and DiscussionResultsIn general, the actuator performed well, offering large degrees of deformation with no signs of mechanical failure in any of the films, actuator members, or compliant joints. Fig. 10 (b) shows a superimposed image of the actuator with the films in their fully wet and dry states during one of the testing cycles.

Fig. 10 (b) shows a superimposed image of the actuator with the films in their fully wet and dry states during one of the testing cycles. This is compared with the targeted design of wet and dry state curvatures as seen in Fig.10 (a). It can be visually observed that while the general shape change was achieved, there was still significant variations from the expected results. To quantify the degree of shape change of the physical experiment(cid:3)(cid:3)(cid:87)(cid:75)e deflections of the actuator were calculated. Variable955deflection of the joint was expected considering that the loading conditions of the joints differed depending upon their location along the cantilever length. However, the stretching, compression, and bending of each joint was too small to accurately measure based solely on the images, and so only the deformation of the entire actuator was measured.

However, the stretching, compression, and bending of each joint was too small to accurately measure based solely on the images, and so only the deformation of the entire actuator was measured. Further steps could be taken in future work to find a reliable means of measuring the joint deflection.aI Dry state II Wet state I',ii ..... 1,iit '-.11· ,Ill .. , -,1 1· -~~., ·r, · ... ...... --=I=t,l ll,llLW'1·1·I I Dry stall'~- ;· - i,;ii- . ; . ~, ~ ~ ,= , ....•. ~.lo' l~,f;:'1 r, , ... :..Wet state II II 11 ~ II II II II•••••••,...-~ -....\ .~ --~\'••1,.. --· ""' . ' ... --:.. ..~ - -.. .... I. '..ii.av1.:rag1..· 7 1.~ mm,/\I 1 , _ __ l 1,u ,•'''--Fig.

--· ""' . ' ... --:.. ..~ - -.. .... I. '..ii.av1.:rag1..· 7 1.~ mm,/\I 1 , _ __ l 1,u ,•'''--Fig. 10 Comparison of (a) Simulated curve of dry and wet state (b) Physical results of dry and wet state curvatureThe average total deformation between the two states of the actuator was calculated to be 71.2mm, measured by changes in height of the cantilevering end of the actuator. This 71.2 mm represents nearly one-third of the total actuator length, which points to the ability of the CM to accommodate a relatively large range of motion. The expected deformation from 2D simulation was 95.6mm, and so evidently the chitosan did not expand to their 12.8 % capacity as expected. It was hypothesized that this was not due to inadequate joint compliance, as the joints already showed much potential for elastic deformation.

It was hypothesized that this was not due to inadequate joint compliance, as the joints already showed much potential for elastic deformation. It is possible that even though the films lose much of their stiffness when saturated, that there was still insufficient driving force to cause significant mechanical strain of the films. One potential workaround would be to implement another tensile element to the assembly that, when added on top of the assembly’s self-weight, could encourage the full elongation of the chitosan films. To determine the actuation time, the video documentation was analyzed to determine at which times the actuators ceased to show discernible changes in curvature during the 20-minute-wetting and 45-minute- drying windows. The time for full actuation from the dry to the wet state averaged on 12.2 minutes, while the reverse process lasted on average 44.2 minutes.

The time for full actuation from the dry to the wet state averaged on 12.2 minutes, while the reverse process lasted on average 44.2 minutes. The faster response time for achieving the wet state was expected due to chitosan’s hydrophilic tendencies. In each testing iteration, the chitosan films (combined weight of 2.75 gm) successfully lifted the actuator’s weight (total weight of 13.5 gm), proving that the films could bear at several times their own weight during drying... .... . -~1-1 ,, ...... ~- ·■ 1/,t,'_~ ,~ ... . ffi .., ••• ••••• •••••••••• ••••••••• ••••••••••illillilaFig. 11 Comparison of curvature over three cycles (a) Dry state (b) Wet stateFig. 11 (a, b) reveals the dry states and wet states for the three testing cycles to understand the repeatability of the shape change. It is clear that the wet state showed excellent consistency among the various experiments, with no distinction in shape seen in cycles 2 and 3.

It is clear that the wet state showed excellent consistency among the various experiments, with no distinction in shape seen in cycles 2 and 3. The dry state, however, possessed some variation, which could possibly be due to uneven drying. It is also wondered if the chitosan film’s drying ability features some slight956957a stepwise sequence of motion of the actuator as it transitioned from dry to wet states was not developed. A robust simulation should be developed, as well as an improved means of measuring the shape changes through technologies such as 3D scanning or motion tracking to accurately compare to the simulation. Finally, while the 3D printed parts were strong enough to handle the loads for this phase of experimentation, the actuator design would need to be substantiated in order to seriously consider this class of actuator for structural applications.

Otherwise, a different 3D printer, material, and/or printing process could be selected, as these decisions presently stemmed from factors of cost and availability.Foreseeable future work lies in scaling up this prototype with tests in real-world environmental conditions, gauging its effectiveness as a 4D compliant structural object. New applications should be conceived, such as developing these structures for rain-responsive shelters or infrastructure. Because the speed of actuation depends on the rate of water absorption by the films, fast-acting movement is possible for climates that experience large amounts of sudden rainfall, such as the tropics. Regions with high amount of relative humidity for sustained periods will also serve to actuate the potential structures to their full capacity, albeit more slowly.

Regions with high amount of relative humidity for sustained periods will also serve to actuate the potential structures to their full capacity, albeit more slowly. However, before long-term outdoor exposure can be considered, the ability of chitosan to survive the onslaught of heat and decomposing microorganisms must be improved with environmental degradation studies. These material science developments should come coupled with AM advancements, which together can hopefully one day enable the creation of fully 3D printed, adaptive biomaterial structures.AcknowledgementsThe authors would like to thank many institutions affiliated with the Singapore University of Technology & Design, including the Fermat Lab, Digital Manufacturing & Design Research Center (DManD), the SUTD-MIT International Design Center (IDC), Architecture & Sustainable Design Pillar (ASD), and the Fabrication Laboratory.References[1][2] [3] [4][5][6][7][8]A. T. Gaynor, N. A. Meisel, C. B.

 Printing of Electro Mechanical SystemsEfrain Aguilera1,2, Jorge Ramos1,3, David Espalin1, Fernando Cedillos1,3, Dan Muse5, Ryan Wicker1,3, Eric MacDonald1,21. W.M. Keck Center for 3D Innovation, University of Texas at El Paso, TX 79968 USA 2. Electrical and Computer Engineering, University of Texas at El Paso, TX 79968 USA 3. Mechanical Engineering, University of Texas at El Paso, TX 79968 USA 4. Materials Science and Engineering, University of Texas at El Paso, TX 79968 USA 5. Printed Device Concepts, LLC, 500 W. Overland, El Paso, TX 79901 USA AbstractRecent research has focused on the fabrication freedom of 3D printing to not only create conceptual models but final end-use products as well. By democratizing the manufacturing process, products will inevitably be fabricated locally and with unit-level customization.

By democratizing the manufacturing process, products will inevitably be fabricated locally and with unit-level customization. For 3D printed end-use products to be profoundly meaningful, the fabrication technologies will be required to enhance the structures with additional features such as electromechanical content. In the last decade, several research groups have reported embedding electronic components and electrical interconnect into 3D printed structures during process interruptions. However, to date there appears to be an absence of fabricated devices with electromechanical functionality in which moving parts with electronic control have been created within a single Additive Manufacturing (AM) build sequence. Moreover, previously reported 3D printed electronics were limited by the use of conductive inks, which serve as electrical interconnect and are commonly known for inadequate conductivity.

Moreover, previously reported 3D printed electronics were limited by the use of conductive inks, which serve as electrical interconnect and are commonly known for inadequate conductivity. This paper describes the fabrication of a high current (>1 amp) electromechanical device through a single hybrid AM build sequence using a uPrint Plus, a relatively low cost 3D. Additionally, a novel integrated process for embedding high performance conductors directly into the thermoplastic FDM substrate is demonstrated.

Additionally, a novel integrated process for embedding high performance conductors directly into the thermoplastic FDM substrate is demonstrated. By avoiding low conductivity inks, high power electromechanical applications are enabled such as 3D printed robotics, UAVs and biomedical devices.Index Terms— Additive Manufacturing; 3D Printed Electronics; 3D Printed Electromechancial devices; hybridmanufacturing; Structural ElectronicsI.INTRODUCTIONAdditive Manufacturing (AM) (more popularly referred to as 3D printing) is a form of manufacturing in which a Computer Aided Design (CAD) model is captured and then subsequently fabricated layer-by-layer. The AM technology used to create the electromechanical part described in this paper is a polymer extrusion-based process – in which a thermoplastic material is extruded through a heated nozzle to provide a robust polymer substrate – in the end fabricating an object that is customized without the requirement for tooling.

Although the process is useful for fabricating mechanical parts such as custom fixtures or molds, the shapes are relegated to simple mechanical objects. Recent research has shifted focus into increasing the functionality of AM- fabricated parts through the integration of electronic components and interconnects into 3D printeddevices [1-6]; however, reports of freeform fabricated electromechanical devices are limited. In [7], AM technology was used to build a simple electrostatic biocompatible actuator with a service life of two to three actuation cycles, this displays a clear interest in increasing functionality of AM parts. Concurrently to and independently of this work, a cleverly designed 3D printed stepper motor has surfaced in pop culture [8], but the fabricated parts required a final assembly process and could have been easily fabricated with traditional manufacturing.

Consequently, the example does not take advantage of having access to individual layers during a build with AM and seems to offer no significant advantage in AM other than the reduction in amortizable cost of the initial units by eliminating expensive tooling. The University of Texas at El Paso (UTEP) is exploring the fabrication of fully electromechanical devices fabricated in a single build sequence that includes an AM technology with process interruptions allowing for the robotic placement of components (microcontrollers, magnets and other electronic control circuits) to create multi- functional 3D printed structures.The advantages of a 3D printed electromechanical device such as a motor – fabricated in one single build sequence as opposed to an assembly process - is the ability to apply the benefits of 3D printing (unit level customization) to not just mechanical parts but also electronic and electromechanical parts within larger customized systems such as a robot or UAV.

When each fabrication can be unit-level customized as required for the final application, the possibility exists to fabricate objects that conform to human anatomy or to provide rapid deployment of a specific UAV system. Other benefits include the ability to provide geometric complexity for free - fabricating structures that cannot be built in any other technology (e.g. internal cavities or meshes). Additionally, by adapting the placement and orientation of the electronics within the final envelope of the design, volumetric efficiency can be dramatically improved.II.Previous WorkRecently, reports in literature describing 3D printing with embedded electronics have been growing.

The combination of Direct Printing (DP) of conductive inks onto 3D printed structures was introduced by Palmer [9] and expanded in Medina [10] and Lopes [11] in which simple circuits were implemented to demonstrate functionality by integrating a dispensing system into a stereolithography (SL) machine using three-dimensional linear stages with a dispensing head. This approach included a demonstration of a simple prototype temperature sensor with nine components including a 555-timer chip. Navarrete [12] describes improvements to using DP on AM substrates by introducing channels into the substrate for the conductive material in order to provide delineation of the electrical lines and allow for the reduction of line width and spacing while simultaneously reducing the possibility of line-to-line shorting. Line pitch was thus determined by the precision of the SL fabrication (e.g. laser beam size) rather than the dispensing process with lower spatial resolution.

Line pitch was thus determined by the precision of the SL fabrication (e.g. laser beam size) rather than the dispensing process with lower spatial resolution. 3D printed electronics can conform to any shape providing a unique advantage over mundane electrical systems and one example of exploiting the 3D design freedom is illustrated in a CubeSat called Trailblazer, which will be launched into low Earth orbit in late 2013. One of the elements of this small satellite is a 3D printed UV curable substrate with embedded electronics, including a microcontroller, gyroscope, crystal, and several passive components (e.g. capacitors and resistors) [13].Though advancements regarding the process of DP electrical interconnects integrated into3D printed structures were described and have contributed to the advancement of low power applications, such as magnetometers or six-sided gaming dice, the research to date described electronics that do not require electric currents exceeding 100mA.

In these low power cases, the conductive ink interconnects are sufficient in terms of conductivity. However, the application described in this paper requires electric currents that can reach 5 amps and possibly higher, which required an appropriately sized or capable current carrying conductor.

However, the application described in this paper requires electric currents that can reach 5 amps and possibly higher, which required an appropriately sized or capable current carrying conductor. A patent pending process developed at UTEP in which solid metal conductors can be embedded into 3D printed thermoplastic structures provides electrical interconnect with sufficiently high current density capabilities for the presented demonstration [13] – similar to the conductivity provided by traditional Printed Circuit Boards (PCB).III.FabricationAlthough many AM technologies exists today that are capable of producing metal, ceramic or plastic substrates, an extrusion-based AM process that deposits thermoplastics was selected to create the 3-Phase DC motor due to ease of access to the build chamber, the low cost of operation, the durability of the thermoplastics and the simplicity of the build process.

The goal of this line of research is to eventually allow consumers to download an electro-mechanical or electronic design from the internet (such as a cell phone or toy car) and 3D print a fully functional device in the comfort of their own home, or alternatively, to allow NASA to manufacture space hardware on the International Space Station or a lunar colony. Plastic extrusion AM uses commercial grade thermoplastics – providing for the fabrication of durable substrates. Additionally a sacrificial support material is used that can be removed in an ultrasonic bath to allow for overhanging substrate features. The AM machine selected to create the 3-Phase motor was a uPrint Plus manufactured by Stratasys (retail price of approximately $20k) and the material used was ABSplus - the material provided by Stratasys [14].

Consequently, this paper demonstrates that electromechanical devices can be fabricated with relative inexpensive printers and that printing multi-functionality is not limited to expensive commercial-grade AM systems. This capability is fertile for exploration and optimization by the Do It Yourself (DIY) community.The main goal of this demonstration, however, is to illustrate the potential for multi- technology fabrication, which is currently under development at the W. M. Keck Center for 3D Innovation. Whereas the demonstration described in this paper was completed (1) with an inexpensive commercial system, and (2) with interruptions introduced for the manual insertion of components and electrical interconnects, the ultimate objective of this course of research is to fabricate a motor in a completely automated fashion with a novel multi-technology AM system (multi3D).

The demonstration illustrates what is possible to fabricate when multiple manufacturing techniques are integrated together. With contemporary AM providing the base fabrication process for a dielectric substrate, a comprehensive manufacturing suite is being integrated seamlessly to include 1) extrusion of a wide variety of robust thermoplastics/metals, 2) micromachining, 3) laser ablation, 4) embedding of wires and fine-pitch meshes submerged within the thermoplastics and 5) component placement. Collectively, the integrated technologies will be capable of fabricating multi-material structures through the integration of a wide variety of manufacturing systems (multi-technology) to provide multi-functional products (consumer wearable electronics, bio- medical devices, defense, space and energy systems, electromechanical devices etc.).IV. Design and Build ProcessThe freedom of design provided by AM technologies allows a unique approach for conformal electromechanical devices.

Design and Build ProcessThe freedom of design provided by AM technologies allows a unique approach for conformal electromechanical devices. For example, the presented electric motor could be fabricated in the form of a conical shape if the intended use is in the nose of an airplane, or alternatively, mounting brackets could be designed onto the motor increasing the freedom of device placement within a large system on a case-by-case basis. This paper describes the process in which a 3-Phase brushless DC motor is constructed in a single build sequence, including a rotor inside a stator separated by two single row ball bearings that serve to maintain alignment between the two structures and allow the rotor to rotate freely within the stator. Designing electromechanical parts requires a more complex design representation using CAD when compared to a simple mechanical form, due to the concomitant mechanical and electronic features.

Designing electromechanical parts requires a more complex design representation using CAD when compared to a simple mechanical form, due to the concomitant mechanical and electronic features. These simultaneous features require both mechanical and electrical CAD methodologies and the confluence of both was completed manually in this example. Future work includes a unified methodology in which 3D printed electronics could be designed with a high degree of freedom component placement and routing within arbitrary forms.The first complication faced when using a uPrint Plus was the lack of available options and control over the rastering and support material. This includes no option for at-will addition or removal of support in the design to be printed. Whereas in some AM system, the user has arbitrary control over the use of support, design rules had to be implemented in the uPrint design to avoid post-processing and other complications associated with the support material.

Eliminating the need for support is important when building internal features since access for support removal may be difficult or impossible. In the case of the presented single motor build, moving internal structures and electronic components were encapsulated during the build process, and consequently, the support cannot be easily removed after fabrication. The removal process involves an ultrasonic bath, where the device is submerged and heated in water for several hours until all the sacrificial material has been dissolved. If water-soluble support material is required to enable internal overhanging structural features and is subsequently removed in an ultrasonic bath after fabrication, electronic components cannot be included in the substrate prior to this post- process step. The minimum angle to avoid the automatic placement of support in the uPrint Plus’s software is 45 degrees as determined through trial and error with the tool.

The minimum angle to avoid the automatic placement of support in the uPrint Plus’s software is 45 degrees as determined through trial and error with the tool. The electromechanical part has to be designed in a way that takes advantage of using a minimum angle of 45 degrees for overhangs in order to avoid required support structures as shown in Figures 1. The form in Figure 1a has an overhang that requires support material, while the shape in Figure 1b is more gradual and can be completed without support material. Removing the requirement for support material has additional advantage of faster fabrication by reducing total deposition time as well as eliminating the requirement for changing between the two materials. For the motor described inthe(a) (b) Figure 1 – Design comparison (a) support needed (b) no support neededthis paper, the only support material used was for the base of the motor.An alternative approach is to use different STL file processing software (e.g.

Insight) that allows for complete control over placement and removal of support material. Normally, the uPrint Plus’s proprietary software automatically places support material to serve temporarily to form a cavity, but for design cases like the one described in this paper, where components are being embedded, the user can identify that the cavity will be filled with a component rather than maintain free space and that once embedded can double as support material. This allows for material to be dispensed over the component and subsequently continue the build. Direct control of deposition of the support material allows the user to evaluate each cavity and decide if a component is present and will act as support.

Direct control of deposition of the support material allows the user to evaluate each cavity and decide if a component is present and will act as support. The lack of additional control provided by the uPrint Plus software led to the implementation of 45 degrees to indirectly avoid support material and allow for cavities inside the device where components can be embedded.A further complication of the 3-Phase motor is that this application requires high electric currents depending on performance and power requirements and was designed to be able to consume up to 10 amps at 12 volts; however similarly sized motors on the market can reach up to 120 amps or more. Initial versions of the UTEP AM motors have been fabricated using electrical interconnects dispensed using a Direct-Write (DW) method for conductive inks - typically consisting of a silver-based ink suspended in a binder.

These types of printed interconnects have very low conductivity due to the restrictions imposed by the plastic substrate on the curing temperatures. Conductive inks generally cannot handle high current densities (>0.02mA/um2) where 150 mA fused a printed line of approximately 250 um width and 10 um thickness – as opposed to 40 gauge wire at 1.5 mA/um2 [15]. When used to connect the electronics and electromagnets of the motor, printed lines were destroyed due to high current densities required and became open circuits – resulting initially in unreliable operation and ultimately in system failure. Therefore the use of conductive inks was not possible. Alternatively, a patent pending process developed at UTEP was employed that embeds solid copper wire directly into thermoplastic substrates by submerging the wire through either ultrasonic welding or joule heating [13].

With 24 to 40 gauge embedded wires (511 microns to 79 microns in diameter, respectively), the current carrying capacity requirements were achieved (32A to 8A, respectively [15]), while simultaneously maintaining high routing densities with center-to-center distances as low as 150 microns. Figure 2 illustrates an 80 micron wire embedded into the thermoplastic substrate through ultrasonic embedding in a conceptual drawing and as a photo of an example fractal antenna.A third challenge of building the motor included the insertion of electronics into partially fabricated structures during interruptions of the extrusion-based AM process - due to the maximum storage temperature of the components. Whereas a stereolithography process provides a build chamber at ambient temperature, the temperature involved in an extrusion system can be damaging to electronic components – particularly at the extruding tip, where temperatures are substantially higher than the rest of the build chamber.

Using a thermocouple the tip temperatureFigure 2 – Thermally embedded wire (a) conceptual depiction of embedding process and (b) in actual example in a 3D printed substrate (fractal antenna)(while extruding) measured 200 degrees Celsius and the chamber temperature was 75 degrees Celsius. The temperature of the extruding tip is sufficient to destroy components - the majority of which have maximum storage temperatures of 150 degrees Celsius or lower. Further studies are necessary to understand the implications of these short durations of applied high temperatures to the functionality of the components. By applying the 45 degree rule, the two main fabrication issues are solved in the context of the uPrint Plus: (a) eliminating internal support material to maintain internal cavities and allow for moving internal structures, and also (b) minimizing the exposure of the components to the extruding tip.

By designing 45-degree “roofs” over every fully enclosed cavity, no support material was necessary, the component becomes fully contained and the components temperatures are reduced. This 45-degree approach, illustrated in several locations of the motor in Figure 3, also has the added benefits of reduced weight in the final structure, material costs and fabrication time.(a) (b) (c)Figure 3 –CAD file (a) rotor (b) stator (c) full drawing, showing cavities with “roofs”, highlighted in yellow are 45 degree features eliminating the use of sacrificial materialAdditive manufacturing also provides the opportunity of adopting a more volumetrically efficient electromechanical part. In this application the electric motor occupied a total of six cubic inches of space, in which included six magnets, nine electromagnets, two ball bearings, and an electronic speed controller, for a total of 18 components.

Each of these components where placed during five pauses within the build - separated into six fabrication segments. The total AM building time of the motor was seven hours and six minutes. Including the duration required for manually embedding components, an additional 39 minutes are required. The five process interrupts are shown in a schematic in Figure 4. The first pause was used to insert the first of two ball bearings to help reduce friction and to provide alignment between the two moving structures: the rotor and the stator. The second process pause was used to insert the magnets inside the cavities of the rotor. Then, the electromagnets were placed and electrically connected in the third process interrupt and treated as pre-manufactured components (similar to the chips and magnets) to avoid the complication of printing the coils within the structure.

The fourth process interrupt was used to insert the second and final ball bearing that completed the two mechanical connections between the subsystems allowing the rotor to maintain concentric position within the stator in conjunction with the first ball bearing. The fifth and last process pause was used to embed the electronic speed controller and connect it to the electromagnets. The last segment encapsulates the electronic speed controller and completes the outer volume of the motor. Once the final segment was finished the motor was removed. The amount of sacrificial material used at the beginning of the build is 4% of the total material and was easily removed from the platform and also from the part.

The amount of sacrificial material used at the beginning of the build is 4% of the total material and was easily removed from the platform and also from the part. Immediately after fabrication, only power and a control signal were required for the motor to operate.Electrical interconnects were completed (soldered by hand) during the third process pause in which nine electromagnets in Wye formation were connected to the electronic speed controller. The three-phase connection from the electromagnets was coupled to the electronic speed controller with enamel wire. Embedding enamel wire into the thermoplastic substrate during the thirdFigure 4 - Pauses in between the buildFigure 5 –Electromagnets embedded Figure 6 – Micro-Welding with YAG laserprocess interrupt provided the additional benefit of maintaining the position of the components during subsequent fabrication. The enamel dielectric avoided shorts between lines as the wires shared a single channel in the substrate.

The enamel dielectric avoided shorts between lines as the wires shared a single channel in the substrate. These channels also provided a flush, planar surface to continue thermoplastic deposition as seen in Figure 5. Connections were then hand soldered on the final pause to the electronic speed controller completing the circuit. As described in Section 2, a system under development (multi3D) will provide the automation of a range of integrated manufacturing technologies such as laser welding systems for a solderless and automated component connection solution. An example of a laser-welded component is shown Figure 6. Implementing a YAG laser micro-welding system from Miyachi Unitek [16] within a custom machine like the Multi3D is the subject of ongoing work for the fully automated fabrication ofFigure 7 - Testing set-upelectronic and electromechanical devices.V.

ResultsAn electronic speed controller (ESC) is used to control the speed of the motor and requires power and a Pulse Width Modulated (PWM) control signal. A microcontroller or a function generator is used to create the PWM signal - the duty cycle of which is proportional to the speed of the motor. Connecting the 3-Phase printed motor to a 12 V power, supply and providing the control signal to the ESC is enough to be able to start and run the 3-Phase DC motor immediately after removal from the building platform. Comparing the 3D printed motor to a commercially- available off-the-shelf 3-Phase brushless DC motor allows for a comparison of how additive manufacturing can compete with traditional manufacturing. The evaluation consisted of comparing torque vs. time and RPM vs. time for an identical start up sequence of power and a PWM signal.

The evaluation consisted of comparing torque vs. time and RPM vs. time for an identical start up sequence of power and a PWM signal. To determine these values, a photodiode and laser system were used to measure the angular velocity and acceleration of the motor as shown in Figure 7. The calculated moment of inertia and measured angular acceleration were used to determine torque.The initial angular acceleration on the 3D printed motor is slower than that of the commercial motor (Figure 8) however after approximately one second, the rate became comparable. The same relationship was seen between the torques of the two motors in Figure 9. Due to initial design constraints of the AM-fabricated device, the motors were different in the number of coils and magnets - with nine electromagnets and six magnets for the AM motor compared to 12 electromagnets and 14 magnets in the commercial motor.

The difference in the number of magnets is the main culprit for the disparity in performance between the motors. The distance between the rotor and stator also played a role in the performance difference where the minimum separation allowed by the software was 1mm. Unfortunately, the difference in architecture does not allow for a direct comparison of the motors, but does provide a general non-Figure 8 – RPM comparisonFigure 9 – Torque comparisonqualitative evaluation. Future work includes fabricating a motor that is identical to a commercial system at least in terms of number of internal components for a more direct comparison.VI. Conclusion/Future WorkAn electromechanical device was fabricated with 3D printing and functionality was demonstrated.

Conclusion/Future WorkAn electromechanical device was fabricated with 3D printing and functionality was demonstrated. Using relatively inexpensive commercially available equipment, printing of electronic and electromechanical devices is possible in remote places such as space or battlefields where no readily available inventory is in stock – or by the DIY community. Inevitably, even consumers will be able to download printable design files representing an electronic device, and subsequently, 3D print the device within their own home.Though manual intervention was used in this example to build an electromechanical device, this system demonstrated the potential of what can be fabricated with a multi-material and multi- technology 3D printer. Current work includes seamlessly integrating several manufacturing technologies into a single system (Multi3D) including multiple polymer extrusion technologies, micro-machining, wire embedding, pick and place and micro-welding.

Once complete, electromechanical systems such as the demonstration piece described in this paper will be fabricated in a single build sequence without manual intervention. Eliminating human involvement will allow for improved manufacturing tolerances and simplicity of fabrication and eventually will enable a “press print” methodology where one can start a fabrication of a CAD design and return to a new functional device the next morning.AcknowledgementsThe research presented here was conducted at The University of Texas at El Paso within the W.M. Keck Center for 3D Innovation (Keck Center). Through funding from the State of Texas0Emerging Technology Fund, the Keck Center recently expanded to over 13,000 sq. ft., housing state-of-the-art facilities and equipment for additive manufacturing processes, materials, and applications.

ft., housing state-of-the-art facilities and equipment for additive manufacturing processes, materials, and applications. Specifically, this research was supported by funding from the State of Texas Emerging Technology Fund as well as University of Texas System Louis Stokes Alliance Minority Participation (LSAMP) under grant NSF grant number HRD-1139929. The authors are grateful to Karla Carmona from the University of Texas at El Paso’s Physics Department and Elaine Maestas for their participation and contribution.References[1] A. Lopes, E. MacDonald, R. Wicker, “Integrating Stereolithography and Direct Print for 3D Structural ElectronicsFabrication”, Journal of Rapid Prototyping, Issue 2, Volume 18 (2012).[2] R. Wicker, E. MacDonald, “Multi-material, multi-technology stereolithography”, Virtual and PhysicalPrototyping, Vol. 7. Issue 3, pp. 181-194.[3] C. Gutierrez, R. Salas, G. Hernandez, D. Muse, R. Olivas, E. MacDonald, M.D. Irwin, R.B. Wicker, K. Church,M.

7. Issue 3, pp. 181-194.[3] C. Gutierrez, R. Salas, G. Hernandez, D. Muse, R. Olivas, E. MacDonald, M.D. Irwin, R.B. Wicker, K. Church,M. Newton and B. Zufelt, “CubeSat Fabrication through Additive Manufacturing and Micro-Dispensing,”Proceedings from the IMAPS Symposium, Nov. 2011.[4] Castillo S., Muse D., Medina F., MacDonald E. and Wicker R. (2009), “Electronics integration in conformalsubstrates fabricated with additive layered manufacturing”, Proceedings of the 20th Annual Solid FreeformFabrication Symposium, University of Texas at Austin, Austin, TX, pp. 730-737.[5] De Nava E., Muse D., Castillo S., Alawneh M., Navarrete M., Lopes A.J., MacDonald E. and Wicker R.B. (2008),“3D off-axis component placement and routing with solid freeform fabrication”, Proceedings from the 19thAnnual Solid Freeform Fabrication Symposium, University of Texas at Austin, August 3-5, pp.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed Paper3D Printing of Passive Microfluidic Flow Mixers Using Triply Period Minimal Surface Microlattice StructuresMazher Iqbal MohammedSchool of Design and Creative Arts, Loughborough University, Loughborough, UK, LE11 3TUAbstractMicrofluidics are miniaturised devices useful for precision fluid handling phases when conducting a range of chemical reactions or biological processes. Such devices operate at micrometre length scales, where laminar flow dominates and so interactions are limited to diffusion between the flowing liquid interfaces unless flow is made turbulent to induce mixing. Passive mixers are desirable for this task as they comprise geometrical features which can be incorporated during the fabrication of such devices.

Passive mixers are desirable for this task as they comprise geometrical features which can be incorporated during the fabrication of such devices. Designs largely remain planar due to traditional microfluidic manufacturing being conducted with 2.5D fabrication processes. Additive Manufacturing now allows for passive mixers to now be realised in true 3D but have seen limited investigation. This study explores the efficacy of several miniaturised Triply Period Minimal Surface micro-lattice structures, formed within microfluidic channels as turbulence inducing structures for increased mixing. We explore several lattice designs and report on their efficacy for mixing reactions conducted during continuous flow conditions.1 IntroductionMicrofluidic devices have proven a versatile tool for a range of fluidic handling processes, which have found usefulness in field ranging across chemistry, biomedical diagnostics, and environmental sampling [1,2].

Interest in microfluidics is due to the length scales of such devices, which enables them to consume low volumes of what can often be expensive/limited reagents, they can allow for multiplex reaction processes and importantly have high surface to volume ratios, resulting in fast reaction development to endpoint as compared to reactions conducted at the macro scale [3].Historically, microfluidic devices are manufactured using processes adopted from the microelectronics industry, in particular photolithography. Manufacturing of microfluidic devices requires the use of a cleanroom and several labour intensive and time-consuming phases to create a single device. Consequently, to overcome these shortcomings a second generation of device development occurred using rapid prototyping techniques such as laser cutting/engraving [4], micro milling [5], and more recently using Additive Manufacturing (AM) processes [6].

AM technology has matured considerably over the last decade, with a growing wealth of different printing systems each capable of utilising a wide range of materials. As such microfluidic devices have been demonstrated utilising Selective Laser Melting (SLM) [7], stereolithography (SLA) [8], Fused Filament Fabrication (FFF) [9] and agarose extrusion printing to form hydrogel channels [10]. Of all these processes SLA AM holds considerable1692potential for microfluidic applications, owing to the availability of transparent material variants, which allow for easy visualisation of chemical/biological processes, prints have low surface roughness, and fabrication can potentially achieve length scales of the order of one to hundreds of microns. As such there is growing interest in SLA for microfluidic applications.The length scales of microfluidic systems cause laminar flow for most flow regimes [11].

As such there is growing interest in SLA for microfluidic applications.The length scales of microfluidic systems cause laminar flow for most flow regimes [11]. The consequence of this is that migrating fluids remain in confined flow streams with very little mixing occurring outside of modest diffusion at the flowing fluid boundaries. To resolve this issue, the use of active or passive mixers can disrupt laminar flow by inducing inertial forces within the bulk of the flowing fluid [12]. While both are effective means of mixing, passive structures are desirable as they require no moving components, reducing design and fabrication complexity and can be manufactured using the same fabrication processes used to form the microfluidic channel structures [13]. Historically, passive mixing structures have generally been designed to be planar in geometry, largely based on the design constraints of the typically employed 2.5D manufacturing processes (e.g.

Historically, passive mixing structures have generally been designed to be planar in geometry, largely based on the design constraints of the typically employed 2.5D manufacturing processes (e.g. photolithography, laser engraving, etc). With the new generation of AM fabricated microfluidic systems, there is now the potential to realise mixing structures which are designed in true 3D within fluidic channels, generating new modes of turbulence. Lattice structures would be ideal candidates for such applications, as they comprise complex geometrical configurations with porosity, which would allow for free migration of a fluid through the bulk of the structure. Of the wide number of possible lattice configurations, Triply Period Minimal Surface (TPMS) geometries have been gathering considerable interest owing to their complex symmetries [14] and geometry which does not require the use of support structures [15].

Recent studies have also begun to explore the uses of TPMS and wider micro-lattice structures in applications for impact absorption [16] and acoustic dampening [17], however none have explored uses in microfluidic applications.In this study we present preliminary work regarding the fabrication and experimental examination of TPMS lattice structures manufactured in situ within microfluidic channels, with the purpose of inducing mixing of two migrating fluids which otherwise would experience laminar flow. We believe this to be the first study which has made use of lattice structure for such an application. As microfluidic devices are systems which geometrically are at micro length scales, we focus on the manufacturability of such systems using SLA fabrication processes, exploring smallest attainable channel size, alongside fabrication dimensional accuracy.

We then demonstrate the capability of creating repeatable microchannels containing lattice structures, alongside a simple and novel lamination fabrication method to seal the devices for use. We also present a universal luer lock design which complements the intended fluidic flow-based application, before assessing the mixing within such structures. Surprisingly, it was found that laminar flow can occur through the bulk of the lattice structure despite the geometrical barriers inducing inertial forces within the flow, and therefore turbulence. However, we demonstrate that through an interplay of lattice size, unit cell type and flow rate of the migrating fluid, we can generate turbulent flow and mixing of two, coloured dye solutions migrating through the lattice structure.

We present our initial findings and hope this to be a catalyst for further inquiry.16932 Methods2.1 Microfluidics designMicrofluidic networks were designed using a combination of Fusion 360 (Autodesk, UK) and nTopology 3.44.4 (nTopology, USA) CAD software. The overall design of a fluid device comprised a double laminate layer. The lower layers were designed using nToplogy and comprised a standard t-channel mixing system, with entry channels 5mm long and a mixing channel which was 30mm long. The various lattice structures were designed to be the height of the channel and placed 2mm into the straight mixing channel of the t-chip. Lattice structures were formed from various Gyroid and Diamond TPMS unit cells. Fusion 360 was used to form the top layer of each device, which seals the channel and has ports for fluidic entry. For testing purposes, microfluidic channels were kept to a fixed cross section, comprising a height and width of 1mm.

For testing purposes, microfluidic channels were kept to a fixed cross section, comprising a height and width of 1mm. The TPMS structures were trialled with varying unit cell sizes of 0.3 to 1mm, which are within the acceptable size range for the proposed channels. For fabrication tests, a unit cell of a given size was projected through a rectangular structure, with a cross section of 1x1 and length 30mm. For fluidic testing a given unit cell was projected through a 5mm long rectangle with a cross section of 1x1mm.2.2 Microfluidic manufacturingFrom preliminary in-house testing (results not shown) it was found that direct AM of closed microfluidic devices with the TPMS structures containing within a channel resulted in blockages which impeded fluid flow until relatively large design geometries (approximately >3mm features).

This was in part due to the enclosed area impeding resin removal within the channel and lattice during fabrication, and consequently resulting in partial curing within the channel/lattice structures. By contrast, printing the microchannel and microlattice structures as open, unsealed structures allowed for uncured resin to migrate out of channel and lattice structures due to gravity, resulting in superior design reproduction. Owing to the improved fabrication reproduction, open structures were considered the preferred approach in this study, however necessitated post sealing of the fluidic chip through lamination.The lower laminate layer containing the lattice and microfluidic channels structures were manufactured using a Direct Light Projection (DLP) based vat polymerisation 3D printing process using a commercially available desktop printing system (LD006, Creality, China).

This system operated such that the printing build platform is inverted and the raises and lowers into a vat of photocurable polymer, which contains a clear window at the bottom for the projection of ultraviolet light from a TFT screen. The final devices are fabricated by the build platform submerging within the resin to a set distance (layer height) from the TFT screen, where the pattern of the microchip layer is projected into the vat by selective lighting of the TFT screens pixels, resulting in polymerisation of the resin onto the build surface. This process is then repeated layer by layer until the complete chip has been fabricated. The system has a manufacturer stated resolution of 0.05mm in the XY and 0.01mm in the Z. All microfluidic chips were manufactured using a commercially available clear acrylate-based photopolymer (Clear Hard Resin, Technology Outlet, UK).

All microfluidic chips were manufactured using a commercially available clear acrylate-based photopolymer (Clear Hard Resin, Technology Outlet, UK). Slicing of a given design and generation of machine printing instructions was achieved using Chitubox open-source software (v1.9.4,1694Chitbubox, China), using a fixed layer height of 30µm, a bottom layer count of 10, layer exposure time of 2s and bottom layer exposure time of 15s.Once a respective microfluidic chip had been fabricated and removed from the printer, it was placed into an agitating cleaning bath (Form Wash, Formlabs, USA) filled with a resin cleaning solution (Resin Cleaner 30, Photocentric, UK) and washed for 15 minutes. Following cleaning, excess residual solution was removed using compressed air, before the device was placed into a UV curing chamber for 30 minutes to allow for the complete curing of resin within the bulk of the device.

Following exposure devices were ready for bonding of the upper laminate layer.The upper laminate layers were fabricated in 2mm thickness acrylic sheets using a commercial CO2 laser cutter (X252, GCC LaserPro, UK). Prior to laser cutting, the acrylic was prepared with double sided tape (3M, UK) without removing the tape cover on the non-contact side and cutting was performed through both the tape and acrylic. The final laminate layers were formed by removal of the tape cover and manually aligning and firmly pressing the cover layer to the microfluidic layer, creating a sealed microfluidic chip. Acrylic was used, as its transparency allowed for clear visualisation within the fluidic chips.

Acrylic was used, as its transparency allowed for clear visualisation within the fluidic chips. To interface the devices with a syringe pump, a universal luer lock entry port structure was designed, and 3D printed in the same clear resin material and adhesive bonded to the opening of the microfluidic chips.2.3 Fluidic flow and imagingFluidic flow in the microfluidic devices was achieved by infusion into the chip using two syringe pumps (NE-1002X, New Era, USA), which allowed for flow rate adjustment from 1µl/min to 600µl/min. To visualise flow, various food dye solutions were used, which allowed sufficient contrast of the flowing fluid front and also for mixing to be assessed based on the proportion of new colour formation from the two input colours.

Flow and mixing of fluids on chip were imaged using a USB microscope camera (AM4917MZT, Dinolite, Taiwan).3 Results3.1 Print Fabrication Resolution TestsMicrofluidic devices are typically at length scales from a single to hundreds of micrometres, which makes their fabrication often quite challenging. For many AM technologies this would mean fabrication near the limits of build resolution. As such it is important to appreciate the typical length scales which are achievable for both fabrication of the microfluidic channels and the internal lattice structures, with lattice features likely to be an order of magnitude smaller. To assess the attainable build resolution, tests were initially conducted to determine the lower build limits with the given 3D printer and resin combination. This was achieved through the design of a test structure, which included various rectangular channel and raised structures, and which can be seen in figures 1 and 2.

This was achieved through the design of a test structure, which included various rectangular channel and raised structures, and which can be seen in figures 1 and 2. In creating these test structures, we were also able to determine any discrepancies between the CAD and fabricated structures. Figure 1 shows graphs of the width and height of the constructed channel for 5 repeated measurements of the channel cross sectional geometry, including ±1 standard deviation error.1695Figure 1: a) Graphs comparing the fabricated Vs designed i) height and ii) width of the channels, alongside bar graphs illustrating the difference between designed and fabricated iii) width and iv) height. b) A microscope image composite of the fabricated channel cross sections.To assess the difference in the horizonal and vertical print resolutions, channels were designed to be of an equal width and height.

It was found that the 3D printer system could repeatably produce channels down to a cross sectional area of 0.3x0.3mm. However, it was found that the resulting channels could often be partially blocked meaning while the resolution was attainable it was not repeatable for the given settings. We therefore concluded a channel cross section of 0.4x0.4mm to be the lower limit of channel fabrication. It was also found that the fabricated widths and heights were consistently smaller than the design geometry, with differences ranging from 0.4 to 7.1% in width and 1.8 to 8.2% in height as compared to the design geometry. We believe this is due to the inherent shrinkage of the polymer which occurs during the photo polymerisation phases of fabrication, and which has been reported in previous studies [18]. It is also noted that the smallest attainable print feature size is considerably larger than the print resolution of the printer.

It is also noted that the smallest attainable print feature size is considerably larger than the print resolution of the printer. This is believed to be due to the relative concentration of the photoinitiator within the resin, which has been found previously to impact crosslinking density, and therefore print feature size [19]. Interestingly, in microfluidic research, it is often desirable to obtain the smallest resolution of channel structure possible due to the reduced1696reagent consumption and increased reaction mixing dynamics. However, this reduction in size is marginal, measured in this study as between 4-11µm, and which reduced as the geometry approach the limits of print resolution. It was also found that the difference between designed and measured channel width/height began to increase as the length scales became larger, increasing the difference up to 115µm at geometries >1.4mm.

This again would be expected due to the cumulative shrinking effect over many more layers at these geometries. Figure 1b) shows images of the fabricated channel cross sections, where it can be seen there is reasonable geometrical accuracy, resulting in well-defined square cross sections. It is noted that at the bottom of the channels, at the intersection of the horizontal base and vertical side walls, that the fabricated geometry becomes more curved which we believe is due to surface tension effects of the resin during photopolymerisation.Figure 2: a) Graphs comparing the fabricated Vs designed i) height and ii) width of the raised square structures, alongside bar graphs illustrating the difference between designed and fabricated iii) width and iv) height.

b) A microscope image composite of the fabricated raised square structure cross sections.A second set of fabrication tests with a raised rectangular structure was conducted to assess the print resolution when fabricating the solid lattice structures within the channels. Figure 2 shows graphs illustrating the width and height of the constructed raised features for1697five repeat measurements of the structure cross sectional geometry. The graphs including errors of ±1 standard deviation.Comparing the results for the channel fabrication it was found that there was greater reproducibility for the width geometry, with all but the structure made at a width of 1.2mm showing deviation of >10µm between designed and fabricated structures. The higher deviation between design and fabricated structure for a width of 1.2mm was consistent between repeat tests and it is unclear why this was occurring.

The higher deviation between design and fabricated structure for a width of 1.2mm was consistent between repeat tests and it is unclear why this was occurring. Ultimately, this geometry was larger than that which we would be utilising for the fabrication of the lattice structures as so the phenomenon was not investigated further. Examining the height of the raised structures, it was found there was significantly higher variability between design and fabrication heights, which reached a maximum of up to a 12.3% reduction. As with the channel fabrication tests this reduction in height can be explained in part due to shrinkage, however the higher percentage would imply a greater shrinkage. Another possible explanation could be due to the nature of the printing process given fabrication occurs with structures on the build platform being raised and lowered into the vat.

Another possible explanation could be due to the nature of the printing process given fabrication occurs with structures on the build platform being raised and lowered into the vat. During this time either surface tension effects between the part and the exposure area or minoring shearing of uncured surface material could explain the reduced height. It is noted that in figure 2b) several of the imaged structures have a distorted upper region to the square structure, which would be consistent with either damage or non-uniform curing. This would therefore support the argument that the height differences are due to shearing or surface tension effects, we hope to examine the cause of this issue in future work.Figure 3: Microscope images of the fabricated micro-lattice structures for a Diamond unit cell size of a) 0.8mm, b) 0.6mm and c) 0.4mm. d) a Graph illustrated the designed and fabricated micro-lattice structures for unit cell size against the resulting pore size.

d) a Graph illustrated the designed and fabricated micro-lattice structures for unit cell size against the resulting pore size. Note: fabrication data only includes for 0.8, 0.7 and 0.6mm unit cell sizes, as below these geometries the lattice structures were found to be closed.16983.2 Micro-lattice Fabrication TestsThe proposed lattice structures in this study are to be printed at micro scale geometries, which would be at the limit of resolution of the printer. To therefore verify the smallest attainable lattice structures and their structural repeatability several fabrication tests were conducted to determine a suitable geometry for flow mixing tests. Figure 3 illustrates the fabrication tests for a Diamond TPMS structure which was designed and fabricated using unit cell sizes of 0.3, 0.4, 0.5, 0.6, 0.7 and 0.8mm.

Figure 3 illustrates the fabrication tests for a Diamond TPMS structure which was designed and fabricated using unit cell sizes of 0.3, 0.4, 0.5, 0.6, 0.7 and 0.8mm. It was found that for unit cell sizes of ≤0.5mm fabricated structures would have closed pores within the scaffold and so the fabrication limit was determined to be a minimum unit cell size of 0.6mm. As it can be seen in the graph in figure 3d), the fabricated pore size was marginally larger than the designed pore size by approximately 8-19µm. It is not entirely clear why the fabricated geometry would be larger than the design geometry given shrinkage effects. This may be related to surface tension effects of the resin within the Diamond lattice structure during fabrication or could be due to marginal over exposure of the structure given the spatial geometry. Further adjustments to the exposure were not explored but will be the subject of future work.

Further adjustments to the exposure were not explored but will be the subject of future work. The limits of design for fabrication revealed a lower pore size of 238±3.6µm, which can be assumed to be the fluidic channel geometry when a fluidic would migrate through such a lattice. Interestingly, this geometry is approximately twice as small as the smallest, repeatably fabricated microfluidic channel.Figure 4: Microscope images of the fabricated micro-lattice structures for a Gyroid unit cell size of a) 0.8mm, b) 0.6mm and c) 0.4mm. d) a Graph illustrated the designed and fabricated micro-lattice structures for unit cell size against the resulting pore size. Note: fabrication data only includes for 0.8, 0.7 and 0.6mm unit cell sizes, as below these geometries the lattice structures were found to be closed.Figure 4 illustrates the fabrication tests for various Gyroid TPMS structures which were again designed and fabricated using unit cell sizes of 0.3, 0.4, 0.5, 0.6, 0.7 and 0.8mm.

As with1699the diamond structures, it was found that for unit cell sizes of ≤0.5mm, fabricated structures would have closed pores within the scaffold and so the fabrication limit was determined to be a minimum unit cell size of 0.6mm. Examining the graph in figure 4d, the fabricated pore size was lower than the designed pore size by approximately 74-111µm, which amounted to a 20- 34% reduction in size compared to the design. The reduction in size between the fabricated and designed pores are consistent with the previous results which showed a shrinkage in the material. However, on this occasion shrinkage is significantly larger than that found during the channel and raised structure tests. Based on our finding we believe this to be geometry related given the differences between Gyroid and Diamond designs, and how these influence the resin residency within the structures due to surface tension effects of the resin within the lattice pores.

We hope to further investigate material shrinkage in future work. The limits of design for fabrication revealed a lower pore size of 180±4.2µm, which again can be assumed to be the fluidic channel geometry when a fluidic would migrate through such a lattice.3.3 Fluidic Mixing testsFollowing confirmation of design for fabrication limits within the SLA system, a trial geometry comprising channel cross section of 1x1mm was selected for flow testing. Given that the minimum cell size was found to be 600μm, a test TPMS lattice unit cell of 1mm was tested to understand the mixing potential of the micro-lattice.Figure 5: a) The experimental set-up for flow testing, comprising two syringe pumps, various silicone tubing and Luer locks to interface with the microfluidic device and a USB microscope to visualise flow. b) A diagram illustrating the laminate layer design of the fluidic chips, alongside the 3D Printed Luer lock adaptors.

b) A diagram illustrating the laminate layer design of the fluidic chips, alongside the 3D Printed Luer lock adaptors. c) A close-up view of the fluidic chip and d) An example of a fluidic chip containing a Diamond lattice structure, where mixing of the red and blue inputs results in purple at the output.1700Figure 5 illustrates the experimental set up for the flow tests, along with an example of a final microfluidic chip fabricated using laminate hybrid manufacturing, combining 3D printing and laser cutting processes. Initially, a t-chip containing no lattice structures was fabricated to test the flow dynamics, where it was found that laminar flow was observed across the examined flow rates between 1µl/min and 600µl/min (results not shown). Given the geometry of the microfluidic channels, a minimum flow rate of 20µl/min was utilised such that flow dynamics would reach a steady state over the course of minutes, allowing for a rapid assessment of changes in flow kinematics.

The flow rate was also kept the same for both flow streams to ensure there was an even volume of each food dye during interactions.Figure 6: a) Images of the steady state flow within the micro-lattice containing microfluidic devices for changes in the flow rate of the input flow channels. b) A graph illustrating the change in the colour at the output of the microfluidic channel, with greater depth of colour change signifying increased mixing.Figure 6a shows the steady state flow streams for various flow rates within a t-chip containing a 5mm length of Diamond micro-scaffold. It was observed that there was disruption of laminar flow for flow rates >60µl/min. Mixing occurs primarily due to either diffusion within the bulk of a fluid, or due to inertial forces within a microfluidic system, as diffusion is limited in1701laminar flow, all mixing is likely occurring due to inertial effects.

It is therefore believed that the inertial forces due to the directional changes in the micro-lattice are responsible for mixing. Interestingly, it was found that the flow channels switched sides within the single channel from the input flow streams, as can be seen for the 20µl/min flow tests. Typically, this would not be observed in such systems. The migration of the fluids to opposite sides of the microfluidic channel may imply that the two flow streams effectively take a zig zag path as they migrate through the micro-lattice and which has a small radius of curvature, thereby not generating substantial inertial forces and mixing. Additionally, the number of repeating unit cells within a length of micro-lattice may also influence the migration. In these tests there were five complete unit cells in the lattice, and so the odd number of cells may cause the migration.

In these tests there were five complete unit cells in the lattice, and so the odd number of cells may cause the migration. To test this assumption a second system was created with a two unit cell containing micro-scaffold and indeed the flow was found to remain unperturbed (results not shown). We believe this to be the first report of geometrical changes within a channel structure allowing for migration of laminar flow paths to new regions of a microfluidic channel.Another interesting observed phenomenon was the preservation of partial flow streams, which were observed at flow rates of ≥200ul/min, which is visible in figure 6a) as the small blue flow stream at the centre of the channel. The dynamics of how this is occurring is less understood but will be the subject of future studies.

The dynamics of how this is occurring is less understood but will be the subject of future studies. Ultimately, it was demonstrated that the micro-lattice structure provides a passive means of generating turbulent flow within a 3D printed microfluidic system, albeit within a specific range of flow rates.Figure 7: Time lapsed images of fluid entering a t-chip containing a) a body centred cubic lattice structure and b) a Gyroid structure.3.4 Laminar Flow within Micro LatticeDuring the study, an unexpected phenomenon was discovered within the micro-lattice structures, namely that for certain configurations laminar flow was found to continue through the bulk of the lattice and beyond the lattice boundary. Figure 7 illustrates two configurations of lattice structure which exhibited laminar flow through the bulk of the structure, with the unit1702cells comprising a Gyroid and body centred cubic geometries.

Laminar flow within the lattice structures was observed at all previously tested flow rates, which was again surprising to observe. The reason for why laminar flow is occurring is uncertain, as the lattice was expected to act as an effective barrier to the directional flow of the fluid, resulting in inertial effects and ultimately mixing, however this is not what is occurring. One possible reason for the laminar flow could be that the bulk flow in the direction of the channel is larger than that which would undergo inertial effects through direction changes of the fluid, resulting in effective viscous drag force acting on the deflected fluids, driving them back into laminar flow. We hope to investigate this assertion in future work.4 ConclusionThis study successfully demonstrates the potential of TPMS micro-lattice structures as passive mixers to disrupt laminar flow within AM fabricated microfluidic channels.

We primarily show that for Diamond unit cells mixing and a range of different flow phenomena are possible. Firstly, it was found that the micro-lattice structure can move the occupancy of a flowing fluidic within a channel at lower flow rates (<200µl/min), resulting in the red and blue dye solution laminar flow streams moving from left to right respectively. This was believed to be due to the preferred fluidic flow path within the micro-lattice being an effective zigzag through the structure with a low radius of curvature for flow directional changes. These low radii of curvature of the pore flow path results in low inertial forces allowing for flow to remain laminar. Secondly, at a critical flow rate (>200ul/min), sufficient inertial forces are generated for fluids migrating through the examined Diamond micro-lattice geometry resulting in mixing.

Secondly, at a critical flow rate (>200ul/min), sufficient inertial forces are generated for fluids migrating through the examined Diamond micro-lattice geometry resulting in mixing. However, for the body centred cubic and Gyroid micro-lattices it was found that laminar flow occurring through the bulk of specific micro-lattice across all tested flow rates. This result was surprising, and the nature of this phenomenon is geometrically driven, implying certain unit cells may not be suitable for mixing of fluids using this approach. Thirdly, the pores produced in the smallest fabricated lattice structures allowed for the creation of fluidic channels which were approximately half the cross section of the smallest attainable microfluidic channel (approximately 180µm).

Beyond microfluidic systems, lattice mediated fluidic exchange is becoming increasing popular for additively manufactured heat exchange products which would be susceptible to laminar flow, thereby limiting heat exchange capacity. Therefore, these findings may help inform suitable unit cell use and wider heat exchanger designs. It is ultimately hoped that these preliminary findings may help advance investigation of fluidic based exchange in a wider range of lattice structures.References[1] Soumyabrata, B., Uchil, A., Kalsang, T., Chakrabarty, S., Ali, M. A., Srisungsitthisunti, P., Mahato, K. K., Surdo, S., and Mazumder, N. -The revolution of PDMS microfluidics in cellular biology - Critical Reviews in Biotechnology, 1-19, 2022.1703[2] Mohammed, M. I. and Desmulliez, M. P. Y. - Autonomous capillary microfluidic system with embedded optics for improved troponin I cardiac biomarker detection - Biosensors and Bioelectronics, vol. 61, pp. 478-484, 2014.[3] Mohammed, M.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference3D PRINTING OF SHAPE CHANGING POLYMER STRUCTURES: DESIGN AND CHARACTERIZATION OF MATERIALSS. N. R. Kantareddy, T. W. Simpson, Z. Ounaies, and M. Frecker Department of Mechanical & Nuclear Engineering The Pennsylvania State University, University Park, PA 16802AbstractAdditive manufacturing (AM) gives engineers unprecedented design and material freedom, providing the ability to 3D print polymer structures that can change shape. Many of these Shape Memory Polymer (SMP) structures require multi-material composites, and different programmed shapes can be achieved by designing and engineering these composites to fold and unfold at different rates.

To enable SMP to have an applications understanding of the relationships between intermediate shapes and the initial and final designed shapes. To accomplish this, we investigated readily available 3D printable polymer materials and their thermo-mechanical characteristics to create multi-member structures. This paper demonstrates a way to generate different temporary geometric profiles on a single 3D printed shape with the same material. This paper also includes insights from thermo-mechanical analysis of the materials to help create multi-member shape-changing geometries using 3D printing.involving shape-changing geometries,itisimportant1. IntroductionStimuli-responsive materials can change their shape, mechanical properties, electrical properties, optical properties, etc., upon application of external stimulus such as an electrical field, pH, light, magnetic field, thermal field, etc. [1].

[1]. Shape memory polymers (SMP) are such stimuli-responsive materials as they can change from a temporary shape to a permanent shape in response to a stimulus. Heat [2], light [3], and chemical solvents [4] are the few commonly used external stimuli to actuate SMPs [5]. Multi-material 3D printing can be used to create soft polymer composite structures with elastomeric matrix reinforced by such SMP fibers [6,7]. These soft polymer composite structures are also known as Printed Active Composites (PAC) in engineering literature [7] due to the printed composite’s ability to respond to the environment and change shape; however, we use the term SMPs for the remainder of this paper to refer to 3D printed shape memory polymer structures.3D printing is a modern manufacturing method to fabricate 3D objects using layer-by-layer material deposition [8]. CAD data of the 3D object is initially sliced to generate the layer-by-layer contour information.

CAD data of the 3D object is initially sliced to generate the layer-by-layer contour information. There are seven 3D printing processes, and our focus in this paper is the polyjet process, which has been shown to fabricate SMPs [9]. Polyjet-based 3D printing deposits a liquid polymer material and then cures the deposited material with UV light [10]. This continuous deposition and curing occurs until the entire layer is filled with the material, and then the process repeats for the remaining consecutive layers. The Objet 260 Connex3 system [11], which is based on the2224polyjet technique, is used to manufacture the SMPs in this paper. Printing these shape- changing structures is also termed as 4D Printing, where the 4th dimension is time.

Printing these shape- changing structures is also termed as 4D Printing, where the 4th dimension is time. Recently, 4D printing has been gaining attention after the concept was introduced in [12].The 4D printing concept is relatively new, and different research groups are trying to establish its foundation with early investigations into design and materials for SMPs. For instance, researchers have demonstrated methods to use 3D printing to print active materials that can be stimulated with water [13] and heat [14]. Different programmable shapes are also demonstrated in the literature [7,14,15]; however, design guidelines that relate the initial shapes with temporary shapes are still required to enable designers to create complex SMP-based structures.

Although current applications of SMPs are limited due to the difficulty in design and manufacturing of these structures, there are many potential opportunities in future technological applications such automotive seating, customized prosthetics, and gas sensors [20,21]; however, realizing these technological applications is contingent upon being able to design and fabricate different geometries that transform into the desired shapes when activated.There are other classes of materials that exhibit shape memory effects, for example, shape memory alloys [16], but SMPs are unique due to their characteristics of being light weight, flexible, soft, biodegradable and biocompatible [17–19]. Many of these polymer structures require manufacturing of polymer composites with embedded fibers made of a material different from that of the polymer matrix.

Many of these polymer structures require manufacturing of polymer composites with embedded fibers made of a material different from that of the polymer matrix. Manufacturing thin polymer fibers is a critical aspect of manufacturing these SMPs, and melt-spinning has been used extensively to create functional fibers in different profiles [22]. However, this process requires manufacturing fibers first and then embedding them in usable designs for applications. This limits the application of SMPs from design and manufacturing aspect due to the constraints in realizing complex shapes for many different applications. Thus, it is important to develop modern manufacturing methods to create SMPs with flexibility in design, length scales, and allowing for customization from one application to other.

Thus, it is important to develop modern manufacturing methods to create SMPs with flexibility in design, length scales, and allowing for customization from one application to other. Combining SMPs with 3D printing enables creation of such complex geometries because 3D printing is not only suitable for intricate functional geometries, but also allows customization with negligible effect on the manufacturing cost [23–25].The goal in this work is to investigate the mechanical performance of 3D printable materials at different temperatures and demonstrate different temporary shapes on a single 3D printed initial shape. The following section describes the design of the SMPs and 3D printable materials used in this study, followed by results from thermo- mechanical tests. The later section discusses preliminary results obtained from experiments to generate different programmed shapes on the same initial shape for two different materials.2.

The later section discusses preliminary results obtained from experiments to generate different programmed shapes on the same initial shape for two different materials.2. Design and MaterialsThe SMP composites in this study are made of two layers: (1) a flexible layer; and (2) a flexible layer with rigid fibers (see initial shape in Figure 1). The top layer does not contain any rigid fibers, and it is made of a flexible material, available in the Objet2225Connex3, called Tangoblack+. The bottom layer is a composite structure comprised of Tangoblack+ as the matrix material and a blend of Tangoblack+ and Veroclear (a rigid polymer available in the Object Connex3) as the fiber material. Both the Tangoblack+ (flexible material) and Veroclear (rigid material) are proprietary materials of STRATASYS [26].

Both the Tangoblack+ (flexible material) and Veroclear (rigid material) are proprietary materials of STRATASYS [26]. Combining rigid (Veroclear) and flexible (Tangoblack+) materials at different proportions within the Objet Connex3 system creates a hybrid material with intermediate properties; therefore, fibers with different strengths can be fabricated. For instance, a mixture of 5% Tangoblack+ and 95% Veroclear is rigid compared to 100% Tangoblack+, and flexible compared to 100% Veroclear.Figure 1: Shows horizontal bilayer composite transforming to a curved shapeThe shape memory behavior of the composite is due to the difference in the elastic moduli and glass transition temperatures of Tangoblack+ (Tg = -10 to -5 oC) and Veroclear (Tg = 52-54 oC). At room temperature, the rigid material is below the glass transition temperature and exhibits a glassy nature; meanwhile, the flexible material is above the glass transition temperature and exhibits a soft rubbery nature.

When these polymer composites are heated to elevated temperatures (above the glass-transition temperature in this case), they can be deformed by applying small external forces. When the temperature is decreased below the glass-transition temperature, the deformed shape remains fixed in the temporary shape of the structure. When the polymer is re-heated to above the glass-transition temperature, then it is free to regain the original shape at much lower rigidity [18,19]. The process of heating, deformation, unloading and reheating is illustrated in Figure 2.For instance, consider the initial shape shown in Figure 1. Upon cooling, the bottom layer has increased in length as the fibers are stretched and fixed, and the top layer is at the initial/original length. As a result, the polymer attains a curved configuration to accommodate the two layers of different lengths (see temporary shape in Figure 1).

As a result, the polymer attains a curved configuration to accommodate the two layers of different lengths (see temporary shape in Figure 1). Thus, a temporary curvature can be programmed in a straight structure using this process. The composite turns soft and flexible upon heating above the Tg of the fiber material, and the composite returns to its original shape. As a result, the composite changes to its initial shape if left unloaded at high-temperatures.2226Figure 2: Example of thermo-mechanical programming to induce folding in a hinge [15]Different temporary shapes can be achieved by changing the internal fiber arrangement within these structures. Three structures with different fiber arrangements are fabricated as illustrated in Figure 3. Blue and green in the figure represent the matrix (Tangoblack+) and fiber (Veroclear) materials, respectively.

Blue and green in the figure represent the matrix (Tangoblack+) and fiber (Veroclear) materials, respectively. In Figure 3(a), the fibers are continuous in the top layer; whereas in Figure 3 (b)-(c), fibers of different lengths are present in both the top and bottom layers. These composites are fabricated with a blend of 5% Tangoblack+ (for the matrix) and 95% Veroclear (for the fibers). The 3D printed parts are subjected to the process shown in Figure 2, and the resulting temporary shapes are shown in Figure 4.Figure 3: Illustration of different fiber arrangements in compositesAs illustrated in Figure 3, if the fibers are placed in the bottom layer, then the temporary shape is curved towards the top layer. Conversely, if the fibers are placed in the top layer, the shape is curved towards the bottom layer.

Conversely, if the fibers are placed in the top layer, the shape is curved towards the bottom layer. To create a ‘sine’ shape, the fibers are placed in alternate fashion: in the top layer on the left and right ends, and in the middle in the bottom layer (see highlight box in Figure 4 (c). The lengths of these fibers are varied (fibers at the left and right ends are shortened and fibers at the middle are made longer) to generate the omega (Ω) shape (see Figure 4 (d)). These structures regain their initial shape when placed in hot water (around 80 °C). The roll, sine, and omega shaped structures unfold back into the same initial rectangular shape.Although the temporary shapes differ substantially, the initial (flat) shape of these structures is identical; therefore, by controlling the lengths and placement of the fibers in the composite, different temporary shapes can be generated with SMPs.

These different shapes can consolidate into a large SMP structure with more complex profiles.2227Figure 4: Few programmed shapes on an initial rectangular shapeIn the remaining sections of this paper, we investigate the temperature-dependent mechanical performance, and a method to generate different shapes using the same initial structure. Section 3 presents the results of Dynamic Mechanical Analysis (DMA) and tensile tests on different compositions of rigid and flexible materials. Section 4 contains experiments with different stretching lengths to demonstrate various temporary shapes on a single 3D printed shape. Section 4 also presents the recovery times for various materials.

Section 4 also presents the recovery times for various materials. Table 1 summarizes the details of these material characterization tests and shape generating experiments.Table 1: List of samples and related characterizationType of characterizationInitial shapeMaterialsRepetitionsOther parametersDMATensile testsSample 1(5% T) to Sample 4 (20%T) 5%, 10%, 15%, 20%, 25%, and 60% of TTemperature sweep from -50 oC to 100 oCUpto 10% strainGenerating different shapesSee Figure 3(a) See Figure 3 (c)5% T5% T and 20% TStretch lengths: 5.5, 6.5 and 7.5 mm Stretch lengths: 5,6,7, and 8Recovery timesSee Figure 11Polymer 1 (RGD 05), Polymer 2 (RGD 20), Polymer 3 (RGD 30), and Polymer 4(FLX 85)222283. Material Characterization of Different Blends of Rigid and Flexible MaterialsWe use DMA to evaluate the gap between the transitions temperatures in the different material compositions of Tangoblack+ and Veroclear.

Polymer materials are viscoelastic in nature, and the individual contribution of viscous and elastic properties varies with the temperature of the polymer; therefore, moduli in polymers can be denoted by two components: (1) storage modulus, representing the elastic nature of the polymer, and (2) loss modulus, representing the viscous nature of the polymer. As the polymer is heated above the glass transition temperature, the storage modulus (or elastic nature) drops while the loss modulus (or viscous nature) increases. At the transition-temperature of the polymer, where the material starts to exhibit a rubbery nature, the storage modulus decreases significantly, and the loss modulus reaches a maximum.

At the transition-temperature of the polymer, where the material starts to exhibit a rubbery nature, the storage modulus decreases significantly, and the loss modulus reaches a maximum. This transition temperature is also indicated by the peak in the tan (δ) curve that represents the measure of viscous to elastic natures [27].Our DMA tests were conducted on TA RSA-G2 [28], and the experiments were performed over the temperature range of -50 °C to 100 °C at a heating rate of 3 °C /min. The oscillation frequency in the experiment was 6.28 rad/s. Figure 5 shows the peaks of tan (δ) curves that represent the corresponding transition temperatures. The temperature gap (see Figure 5) between the transition temperatures of these materials is narrow (10 °C); therefore, precise thermal stimulation is required to exploit this difference in transition temperatures to actuate folding structures comprised of hinges with different materials.

This precise temperature management is difficult if the hinges are not sufficiently far apart, because thermal actuation of one hinge can indirectly actuate a nearby hinge due to heat transfer.Figure 5: Results of the temperature sweep test on four different material compositions of Tangoblack+ (T) and Veroclear (V); sample 1(5%T) to sample 4 (20%T) are in the increasing order of flexible natureFor many polymer materials, the tensile modulus of the material is also dependent on the strain in the material; therefore, tensile testing is performed to capture the effect of strain on tensile modulus for different material compositions. The tensile test results shown in Figure 6 for the material samples show a general trend that materials that are more flexible have lower modulus for the same strain percent.

The tensile test results shown in Figure 6 for the material samples show a general trend that materials that are more flexible have lower modulus for the same strain percent. Material sample with 10% flexible material deviates from the observed trend at higher strains; however, the trend in tensile modulus at lower strains (less than 1.5%) is still followed. This deviation in2229pattern may be due to manufacturing defects such as lack of proper curing, material non- uniformity, contamination with support material, induced-physical defects, etc.Figure 6: Tensile modulus versus strain curve for different compositions of Tangoblack+ and Veroclear: 5%, 10%, 15%, 20%, 25%, and 60% of flexible material (Tangoblack+)4.

Experiments to Program Different Temporary Shapes in Same Initial ShapeOne way of programming different temporary shapes into SMPs is by leveraging the difference in thermo-mechanical performance of different material compositions; unfortunately, it is difficult to achieve the precise thermal stimulation required to exploit the narrow temperature gap. Another method of achieving different temporary shapes is by controlled deformation of the material during the stretching/deformation phase (see Figure 2). By inducing different deformation lengths, different temporary geometric profiles can be obtained. The experimental setup used in this work to stretch the test specimen by pre-determined values is shown in Figure 7. The setup consists of an assembly of fixed and movable linear stages.

The experimental setup used in this work to stretch the test specimen by pre-determined values is shown in Figure 7. The setup consists of an assembly of fixed and movable linear stages. The gauges attached to these stages can measure the deformation lengths at any moment.A rectangular test specimen is tested by stretching (i.e., deforming) it to different lengths: 5.5 mm, 6.5 mm, and 7.5 mm. The test specimens are subjected to the programming and actuation process illustrated in Figure 2. Once the temperature of each specimen is above the glass transition temperature for the material, then the material can be stretched by turning the knob of the gauge on the movable stage in Figure 7. The deformed/stretched specimen is cooled to room temperature to fix the temporary shape. Three different profiles are programmed on the same initial shape as shown in Figure 8.

The deformed/stretched specimen is cooled to room temperature to fix the temporary shape. Three different profiles are programmed on the same initial shape as shown in Figure 8. This method of inducing measured-deformation at slow rate also prevents the mechanical failure of the material.2230Figure 7: Experimental setup to induce measured deformation in the materialFigure 8: Different programmed shapes obtained by different induced deformation lengths (ΔL): 5.5 mm, 6.5 mm, and 7.5 mmSamples made of two different material compositions (i.e., 95 % rigid and 80% rigid compositions) are subjected to four different elongation lengths to generate different temporary shapes. The resulting temporary shapes along with geometric dimensions are shown in Figure 9. R is the radius of curvature calculated at the center of the temporary shape. For this sample (initial length of 100 mm), decreasing the elongation lengths from 8 mm to 5 mm significantly affected the temporary shape.

For this sample (initial length of 100 mm), decreasing the elongation lengths from 8 mm to 5 mm significantly affected the temporary shape. This procedure demonstrates the ability to induce different geometric profiles by inducing different deformation lengths; however, the repeatability of the programmed profiles is still unknown. To investigate the repeatability, multiple specimens with same material can be tested with the same deformation lengths. Understanding this relationship between material, deformation length, and temporary shapes can help create a mapping between different programmable shapes and initial shapes.The plot in Figure 10 shows the radius of curvature versus the elongation length. We see that a 35% change in radius of curvature can be achieved with a change in elongation length of 20%.

We see that a 35% change in radius of curvature can be achieved with a change in elongation length of 20%. Although the difference in the radius of curvature for two different compositions at same elongation is less, the composition with the higher percentage (95%) of rigid material always yields a lower radius of curvature. This pattern demonstrates that higher radius of curvature can be achieved by using more rigid materials while the stretching length remains same.

This pattern demonstrates that higher radius of curvature can be achieved by using more rigid materials while the stretching length remains same. These preliminary results also demonstrate the ability to achieve different temporary geometric profiles by varying material composition.2231Figure 9: Different temporary shapes programmed on two different materials by stretching at various lengths (8 mm, 7 mm, 6 mm and 5 mm); R is the radius of curvature at center of the shapeFigure 10: Effect of elongation length on the radius of curvature for two differentmaterialsRecovery times of the hinge structures (see Figure 11) made of different compositions are also calculated. Recovery time is the time taken by the sample to shift from a programmed shape to its initial shape upon application of sufficient external heat. For this geometry, the folding angle between the programmed and initial configurations is around 73.5°.

For this geometry, the folding angle between the programmed and initial configurations is around 73.5°. Table 2 presents the recovery times for four different polymer samples. Note that the recovery time is almost doubled when the percentage of flexible material in the composition is increased from 5% to 15%. This difference in recovery time can be useful to create shape-changing structures consisting of members actuating at different rates, but the recovery times are significantly affected by the operating temperatures.2232Figure 11: Shows initial and programmed shape of a two-member hinge; Θ is theangle of foldingTable 2: Recovery times for hinges made of four different materials (blends of Tangoblack+ and Veroclear): Polymer 1 (RGD 05), Polymer 2 (RGD 20), Polymer 3 (RGD 30), and Polymer 4(FLX 85)Material Polymer 1 Polymer 2 Polymer 3 Polymer 4Recovery Time (s) 3.5 5 7 -5.

Closing Remarks & Future WorkThe ability of 3D printing to fabricate shape memory polymers (SMPs) has already been demonstrated, but it remains unclear how to design and actuate initial shapes to obtain different temporary shapes. Different factors such as fiber arrangement, material, deformation length, operation temperature, etc., are critical to achieve the desired temporary shapes. Investigating the thermo-mechanical response of different materials can improve our understanding of how to design and choose materials for initial shapes to achieve a desired temporary shape. First, the difference in transition temperatures of different blends of materials used in making fibers in a polymer matrix structure is investigated. The results show that this gap is narrow and requires precise stimulation of the material to leverage the difference in transition temperature.

The results show that this gap is narrow and requires precise stimulation of the material to leverage the difference in transition temperature. Tensile test results of the material samples show a general trend of changing material properties with change in material composition at higher strains. This difference in material properties can be explored to create range of radius of curvatures by changing material composition of the structure.A method to generate different temporary shapes on a single 3D printed structure by varying the length of stretching (5, 6, 7 and 8 mm) is also investigated. A setup consisting of two linear stages is used to generate temporary shapes of different curvatures. Results demonstrate a potential to generate different shapes by varying material composition and the amount of induced-deformation (stretch length).

Results demonstrate a potential to generate different shapes by varying material composition and the amount of induced-deformation (stretch length). Higher deformation rates are observed to damage/break the sample; therefore, this method of inducing measured-deformation prevents the obvious mechanical failure of the sample. It is also observed that recovery time of temporary shapes significantly depends on the operating temperature.2233Preliminary results in this work demonstrate the ability to achieve different temporary geometric-profiles by deforming the same material at different lengths, illustrating the impact of design decisions on SMPs and their actuation. In continuation of this work, repeatability of these geometric-profiles will be investigated.

In continuation of this work, repeatability of these geometric-profiles will be investigated. Additionally, the relationship between materials, deformation length, and temporary shapes will be thoroughly explored to provide engineers with design guidelines that are required to create the complex SMP-based structures.AcknowledgmentsWe gratefully acknowledge the support of the National Science Foundation EFRI Grant No. 1240459 and the Air Force Office of Scientific Research. Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the views of the National Science Foundation.Bibliography[1] Meng, H., and Li, G., 2013, “A review of stimuli-responsive shape memory[2][3]polymer composites,” Polymer, Vol. 54, No. 9, pp. 2199–2221.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed Paper3D Printing Technology Insertion: Sociotechnical Barriers to AdoptionShawn H. Chang and Bryan R. Moser*System Design and Management, Massachusetts Institute of Technology, Cambridge, MA USAAbstractSince the initial development of three dimensional printing (3DP) in the 1980s, companies have relentlessly researched for applications of the technology. The potential benefit is large, beginning with improved cost and schedule to manufacture plastic and metal articles. As such, governments and industry from advanced economies continue to invest heavily to accelerate 3DP adoption.

As such, governments and industry from advanced economies continue to invest heavily to accelerate 3DP adoption. Amid advancements in the pillars of three dimensional printing – the technology, material, and software – practitioners across industries are steadily deploying 3DP in product development, prototyping, and small scale production of parts and products. However, a large gap remains between promise and the reality of larger scale adoption. The potential benefits, risks, and specific steps to adopt and realize the benefits are not clearly understood, resulting in overly zealous (at risk) or overly cautious (opportunity avoided) approaches to 3DP adoption. Traditional manufacturers rely on decades of know-how in manufacturing practices across a large portfolio of parts, making first steps on a path to adopt new processes more challenging.This paper identifies the variables that complicate or impair judgement when considering the adoption of 3DP.

A systematic approach to evaluate 3DP adoption across a portfolio is needed. A methodology is proposed to analyze the relative value of 3DP at the part and product system level for prototyping and production. The outcome is a framework that combines part-level feasibility with systemic benefit of cost and schedule improvements as prototyping and production alternatives. In building this framework and in interviews with experienced manufacturers, several key insights were gained. Part by part consideration of 3DP feasibility is daunting, while adoption requires readiness not only of 3DP technology but also the receiving systems and organization.

Part by part consideration of 3DP feasibility is daunting, while adoption requires readiness not only of 3DP technology but also the receiving systems and organization. By viewing 3DP insertion as a sociotechnical system implementing the changes, attention is drawn to the tacit knowledge of critical characteristics in existing manufacturing processes, design for manufacturing decisions embedded in existing part assemblies, the pre-processing and post-processing capabilities available to shift 3DP feasibilities, and the alignment of organizational learning across parts.Keywords— additive manufacturing, 3D printing, technology insertion, sociotechnical systems1960IntroductionFor decades the additive manufacturing industry has made significant investment and progress in advancing three dimensional printing (3DP) technologies. Starting from stereolithography in 1986, today there are numerous competing technologies for fabrication ranging from thermoplastics to metal.

Starting from stereolithography in 1986, today there are numerous competing technologies for fabrication ranging from thermoplastics to metal. Their common characteristic is extraction of shape information from three dimensional product models and manufacture by layering the materials in cross sections to produce a final shape. One of the most distinguishing features of 3DP is the ability to create “high geometric freedom” shapes, thus holding “the potential to revolutionize product design”. [1]The applications, from defense to consumer electronics to biomedical, of additive manufacturing have increased significantly in recent years. However additive manufacturing today accounts for less than one percent of all manufactured products in the United States.[2]1.1. Categories and definitions of 3D printing technologiesThere are many categories of additive manufacturing technologies.

Categories and definitions of 3D printing technologiesThere are many categories of additive manufacturing technologies. The ASTM F42 Committee defines additive manufacturing as “a process of joining materials to make objects from 3D model data”. [3] According to the ASTM, the seven primary categories of additive manufacturing processes include binder jetting, directed energy deposition, material extrusion, material jetting, powder bed fusion, sheet lamination, and vat photopolymerization.1.2. Recent industry examplesThe aerospace industry has been an advocate of 3D printing technology and demonstrated significant interests in advancement of the 3D printing technology (including material) and its applications. In recent years, the combination of technology maturity and experience building have resulted in the proven commercial applications of 3D printing for prototype and production intent usage in the industry.GKN TiPOW Research.

GKN Aerospace commences a 3-year collaborative research program to develop titanium powder optimized for AM of aerospace components. The TiPOW (Titanium Powder for net-shape component manufacture) program is backed by UK’s Aerospace Technology Institute and Innovate UK. The program’s objectives are to minimize cost while satisfying the rigorous standards and explore the reuse and recycling of titanium material. GKN Aerospace will supply the turbo exhaust case and compressor intermediate case for Pratt & Whitney.[4]GE T25 Sensor. AA approves the first AM part, GE T25 Sensor, to fly inside GE’s commercial jet engine. The GE T25 sensor houses the compressor inlet temperature sensor in the GE90-94B engines (used on Boeing 777). The part will be produced with cobalt-chrome alloy. GE is currently testing AM-produced fuel nozzles for the next generation LEAP engine and developing AM- produced fuel nozzles for the GE9X (the largest jet engine ever built). [5]Airbus A350XWB.

GE is currently testing AM-produced fuel nozzles for the next generation LEAP engine and developing AM- produced fuel nozzles for the GE9X (the largest jet engine ever built). [5]Airbus A350XWB. Airbus and Stratasys joined efforts in 2013 to develop and certify 3D printing for use in A350 XWB. Over 1,000 flight parts for the A350 XWB were produced using ULTEM™ 9085 resin using FDM technology. The ULTEM™ 9085 resin is certified to an Airbus1961material spec and is flame, smoke, and toxicity (FST) compliant for interior aircraft applications. [6]MTU Aero Boroscope Bosses. MTU Aero mass-produces nickel alloy borescope bosses for the PW1100G-JM geared turbofan (GTF) engine. The borescope bosses are part of the turbine housing for the A320 neo’s GTF engine and allows access to the inspection of the blades using a borescope. MTU plans on building 16 parts per build, up to 2,000 parts/year. MTU and EOS are collaborating to optimize the surface finish of the component.

MTU plans on building 16 parts per build, up to 2,000 parts/year. MTU and EOS are collaborating to optimize the surface finish of the component. [7]Rolls Royce Trent XWB. Rolls Royce uses additive manufacturing to produce the largest 3D printed civil aero engine part – Trent XWB bearing. The bearing is made from titanium and has a diameter of 1.5 meters. It is comprised of “48 airfoil-shaped vane components” each individually produced using EBM technology. Rolls Royce has used AM for 5 years to repair components but is now using AM to make large components from scratch. The bearing will be used in flight for the first time in 2015. [8], [9]1.3. Related WorkStudies to date have been primarily focused on technology readiness or cost effectiveness as the barriers to technology adoption. [2], [10], [11] Other researches have focused on the studying the dynamics of technology adoption in the context of network externalities and assessment of viability.

[2], [10], [11] Other researches have focused on the studying the dynamics of technology adoption in the context of network externalities and assessment of viability. [12], [13] Our research acknowledges those factors and introduces a sociotechnical component that is endogenous to the technology adoption cycle.Farzin describes an approach to calculating the optimal timing to adopt a technology. Leveraging dynamic programming framework, a firm can model the ideal timing to employ a technology especially when the decision is irreversible and uncertainties exists with the maturity and benefit of the technology adoption.[14] Disparities in technology adoption across countries have also been studied. An explanation of this disparity is dependent on the level of “world knowledge” and a country’s inherent barriers to technology adoption.[15], [16]The cost drivers and cost models of 3D printing has been widely researched and published.

Lindemann compared the four main cost drivers of additive manufacturing: build rate, machine utilization, material cost, and machine investment cost.[17] Hopkinson outlined the cost breakouts of a producing a part using different 3D printing processes and presented a cost model based on the calculated average cost per part.[18] Ruffo et al. introduced an activity based approach to cost modeling, with consideration to indirect costs, for selective laser sintering.[19]The interplay during product development between foreground product designs and background design intent have been well documented.

Nonaka characterizes the knowledge management cycle during product development as activity that forms both tacit and explicit knowledge of the product system.[20] Kimura applied this view of deep product development knowledge to CAD systems and the development of production molds.[21], [22] James assessed measure of product complexity in considering the technology insertion into a new variant (geared) of an aerospace turbofan.[23]1962Problem Statement2.1. Product Development as a Sociotechnical SystemStudies on three dimensional printing feasibility have been focused primarily on the technology readiness, part value (cost and time), and process planning (timing, externalities, etc.). Our research focus is on the endogenous sociotechnical dynamics that affect the technology adoption cycle, from which we investigate the attributes that affect the part design, prototype, tooling, and production of parts.

Our objective is to expose the technology adoption risks and benefits at a both local (part) and systematic (product organization) scale as they influence product development choices.Based on several interviews with manufacturing process experts [24], [25], the authors focus was drawn to the barriers to adoption beyond value analysis. The industry experts emphasized that the resultant cost, quality, and schedule of a manufactured part in many cases is the result not simply of an automated process, but instead a combination of technology, process, and the tacit knowledge which aligns the materials, the designs, and the manufacturing environment. An existing part as- designed and as-manufactured embeds a legacy of previous choices in the part itself, and also the pre-processing, material choices, the manufacturing environment, the role of quality during production, post processing, and verification and validation.

This history of choices as background knowledge is both tacit and explicit. Even if a traditional manufacturing technology (e.g. machining) is replaced by 3DP, the remainder of elements in the total engineering and manufacturing process may be still aligned with the previous process. Early failures due to such misalignment may in turn lead to an adoption averse organization.This consideration of the total technical and social systems for product development treats the product part – and its actual cost, scope, quality, and timeliness – as an emergent characteristic of engineering as a sociotechnical system. [26] In this way insertion or displacement of just the part, or a process component, may lead to unanticipated consequences, frustrations, and risks.

[26] In this way insertion or displacement of just the part, or a process component, may lead to unanticipated consequences, frustrations, and risks. Likewise, by considering the broader design of the product development system when inserting 3DP, these frustrations and risks may be avoided.The problem facing a manufacturer considering displacement of existing processes with 3DP may be daunting if the analysis requires a part by part comparison of the multitude of evolving 3DP tools and materials to existing processes.

As shown in Figure 1, the need is for a systematic method to filter and prioritize parts in the context of product and business value, to capture and evolve knowledge about 3DP through this process, and to accelerate learning across the organization by allowing association of insights across the portfolio of parts.1963Figure 1: Challenge to Filter, Prioritize, & Learn SystemicallyApproachFor many companies their current production techniques are backed by years of know-how leading to high quality, reliable parts/ systems. Know-how for new engineering and production methods takes time to develop.

Know-how for new engineering and production methods takes time to develop. In this paper we ask:How to adopt new 3DP product development and production methods at an optimized pace while minimizing risk of poor quality and loss of know-how?Recognizing this key question, the approach should support existing product and process expertsin the field so they can:More rapidly evaluate and prioritize potential application of additive manufacturing to a specific product systemMore rapidly capture lesson learned and insights from this evaluative process, linked (tagged) to product characteristics, process, and additive manufacturing optionsMore efficiently share adoption rules and insights across product systems and business unitsProposed Method4.1. Overview of the MethodThe framework for accelerated adoption of 3DP begins with mapping of major knowledge areas needed to evaluate candidate parts.

Overview of the MethodThe framework for accelerated adoption of 3DP begins with mapping of major knowledge areas needed to evaluate candidate parts. Figure 2 shows product knowledge (P), 3DP tools and material knowledge (T), tacit knowledge of product system engineering, production, and quality (K), and how these combine across an assembly as compared to an alternative baseline manufacturing process. Beginning on the left of Figure 2, an existing product system is represented in a bill of materials (BOM), often containing tens of thousands of parts.

Beginning on the left of Figure 2, an existing product system is represented in a bill of materials (BOM), often containing tens of thousands of parts. The challenge is to search amongst these part candidates and prioritize those that are both feasible for 3DP and valuable to the firm when considering cost, schedule, and product quality at the volumes (and margins) for the business.1964Figure 2: Product, Technology, and Knowledge Combine to Filter & PrioritizeMoving from left to right, from a BOM/parts list to Value, requires attention from various teams across the organization, especially when considering tacit knowledge of part unique production requirements. In many case those teams with deep subject matter knowledge of partwise engineering and production at quality are not the same teams with deep knowledge of market facing volumes, margins, and customer utilities.4.2.

Definition of TermsBelow is a list of terms and corresponding definitions that are used in our methodology.The first set of variables refer to the product system being prototyped and produced.p pn pB pe pmPart (BOM)the nth partbaseline part as generated by existing manufacturing processengineering prototype part(mfg.) production intent partSProduct systemThe next three variables refer to the options for manufacturing the part.TAdditive manufacturing technology; tools and materialBBaseline – the existing (legacy) manufacturing processKBackground knowledge; know-howThe terms below are for evaluating the production optionsF()FeasibilityV()ValuedDurationcCostqQuality19654.3. Outline of MethodologyThe framework begins with the filtering of the Bill of Materials/parts list (BOM) for parts both insourced and outsourced.

Outline of MethodologyThe framework begins with the filtering of the Bill of Materials/parts list (BOM) for parts both insourced and outsourced. Outsourced parts will not be considered and insourced parts will further undergo a part intent filter to classify the parts into prototype intent and manufacture intent.The respective part attributes for prototype and manufacture intents are mapped against the 3DP Technology (T), both tools and materials, for feasibility analysis. Attributes that affect prototyping and production of parts can include the complexity of the part geometry and topology, dimensions, tolerances (dimensional and physical properties), material type and diversity, intended use and environment, appearance (surface finish and aesthetics), strength, certifications, and other relevant elements. In prototyping, concept validation is often the focus and rapid prototyping techniques typically help reduce the product development cycle (time and cost).

In prototyping, concept validation is often the focus and rapid prototyping techniques typically help reduce the product development cycle (time and cost). However, this does not necessarily translate to manufacturability or large scale manufacture as the tools used in prototyping are sometimes not employed in production. This anomaly often leads to well-known risks: namely manufacturability and feasibility. The manufacturability addresses whether a part can be produced to production intent specifications and the feasibility considers the time and cost, which includes sustainability and the yield rate.the dimensional/geometric, accuracy/tolerance, material, mechanical properties, and quality dimensions of the part and the selected 3DP Technology to learn what attributes are within and outside of the limits for the part (p) and 3DP Technology (T).

The feasibility analysis, F(p,T), relies upon part, tool, and material attributes as shown in Figure 3 below.As shownin Figure 3 below,the analysis comparesSome combination of attributes, for example the part dimension and machine capabilities, may quickly filter out the part for consideration. However, some combinations may be nearly feasible yet may be compensated by background knowledge (K), discussed in more detail in the next section. Rather than simply a yes/ no filtration, the analysis exposes the basis of feasibility including risks, with an emphasis on uncertainties to decrease and/or readiness to increase across the p, T, K set.1966Figure 3: Part, Tool, Material Combination Feasibility4.4. Visualization by Feasibility and Valuee) may be deemed slightly less feasible but has more systemic value than the baseline part.

Visualization by Feasibility and Valuee) may be deemed slightly less feasible but has more systemic value than the baseline part. That same part for product intent m) may have more uncertainties in feasibility and systemic value compared to both the prototype (p1 intent and baseline parts. The baseline part may be highly feasible but with somewhat limited systemic value. The visualization helps to prioritize and differentiate the parts for optimal 3DP insertion decision making.In an example shown below in Figure 4, a prototype intent part (p1Figure 4: Comparison of Feasilibty & Systemic Value of 3DP compared to BaselineAn aspect emphasized in the feasibility analysis F(p,T), is the consideration of the background knowledge (K, see Figure 5 below.) The background knowledge (K) may indicate what parts are feasible under current 3DP Technology (T) and tacit know-how such as post-processing capability.

The inclusion of background knowledge (K) also encourages the exploration of the systemic value of the assessed part compare to those of the baseline.19674.5. Iterative Exploration As Basis To Expose And Evolve Tacit KnowledgeFigure 5:Framework with Parts, 3DP Tools and Materials, and Process KnowledgeTaken together, a more detailed view of the method is shown in Figure 5 above. As the large set of total BOM parts is filtered and prioritized, the steps include the ability of experts to review, comment, and in some cases adjust the analysis. At first, with few parts and product systems reviewed, the method will be helpful but still limited to the source knowledge on 3DP technology provided. However, as more and more parts across a portfolio of product systems and 3DP tools and materials are analyzed, then similar parts, technology, and issues (p,T,K) will be linked across the organization.

In particular, as new 3DP tools and materials emerge, this structured knowledge approach can allow an accelerating consideration of adoption across most closely related parts in the organization.Example of Framework Usage5.1. Bill of Material for an Engine Part AssemblyThis example starts with a high level bill of material for a turbine engine part assembly.

Bill of Material for an Engine Part AssemblyThis example starts with a high level bill of material for a turbine engine part assembly. A turbine engine can have over 60,000 unique parts in its Bill of Materials (BOM); in this example we will explore a subset of those parts, focusing on the compressor and combustion sections of the engine.Key parts that make up those sections include the compressor vane ring assembly, compressor blade assembly, combustor blade assembly, shaft, bearings, seal rings and sleeves, expander springs, spline spacers and rings, bearings, couplings, and inlet temperature sensors.1968Figure 6: Typical Jet Engine Diagram[27]The initial step is to filter the compressor and combustion subassembly parts for insource and outsource manufacture. Only insourced parts will undergo the feasibility analysis F(p,T) as the know-hows for outsourced parts are typically exogenous in this framework.

Only insourced parts will undergo the feasibility analysis F(p,T) as the know-hows for outsourced parts are typically exogenous in this framework. Of the key compressor and combustion section parts we assume the bearings, seal rings and sleeves, expander springs, spline spacers and rings, bearings, and couplings are outsourced, while the compressor vane ring assembly, compressor blade assembly, combustor blade assembly, and fuel nozzle are insourced.5.2. Part Usage Intent: Prototype vs. ProductionParts during the design and product development cycle are often instantiated more than once. The next step involves filtering the insource parts into prototype intent (pe) and production intent (pm). Each part will be evaluated based on its part attributes for the usage intent.

The next step involves filtering the insource parts into prototype intent (pe) and production intent (pm). Each part will be evaluated based on its part attributes for the usage intent. For example, early stage prototype intent parts, such as a look-alike prototype, may prioritize on dimensional or “fit” validation and therefore often have different part attributes compared to those of a production intent part. In this scenario, the feasibility analysis should be designed to meet the prototyping requirements and not the exhaustive requirements of those for the final article. However, if the objective was a feasibility analysis for a production intent part, then it will likely be subject to the same requirements, assuming the proper test protocols are known for that 3D technology and material.5.3. Feasibility Analysis for 3D PrintingOnce the parts are filtered by usage intent (pe or pm), their respective part attributes are compared to the 3DP Tech attributes.

Feasibility Analysis for 3D PrintingOnce the parts are filtered by usage intent (pe or pm), their respective part attributes are compared to the 3DP Tech attributes. The 3DP Tech attributes are a summary of the 3DP Technology (T), which includes the available tools and materials for 3D printing. printing tools and materials, denoted as T and part attributes. The 3DP Technology (T) example is the Arcam Q20 industrial metal 3D printer, which employs Electron Beam Melting (EBM) process and is compatible with titanium and cobalt materials. The specific known compatible materials are the Arcam ASTM F75 CoCr (Cobalt-Chromium) Alloy and the Arcam Ti6Al4V (Titanium) Alloy. The build envelope for the Arcam Q20 machine is 350mm x 350mm x 380mm in the x, y, and z dimensions.For part attributes, either pe or pm, it will be existing part specifications with no redesign.

The build envelope for the Arcam Q20 machine is 350mm x 350mm x 380mm in the x, y, and z dimensions.For part attributes, either pe or pm, it will be existing part specifications with no redesign. To start, in the compressor vane assembly, it consists of the compressor ring of over 1200mm in diameter and compressor vane stator blades with common dimensions of 155mm x 40 mm x 40 mm and usually made from titanium. The compressor blade assembly contains the compressor blades of similar size to the vane stator blades and the compressor disc of over 600mm in diameter and both made from titanium. The combustor blade assembly contains the combustor disc that is1969over 600mm in diameter and the blades are also similar in size to the vane stator blades, and with cooling holes for film cooling. The combustor assembly is commonly made from nickel superalloy and sometimes the blades are made from ceramic to withstand the extreme operating conditions.

The combustor assembly is commonly made from nickel superalloy and sometimes the blades are made from ceramic to withstand the extreme operating conditions. The fuel nozzle is commonly produced from cobalt alloy.Knowing the 3DP Tech. attributes (T) and part attributes (pe, pm), the next step would be feasibility analysis of technology and part or the F(p,T) in the framework. The compressor vane ring, compressor disc, and combustion disc parts are too large to be 3D printed in one piece by the Arcam Q20 printer. Furthermore, the combustor disc is made from a material that is incompatible with the printer. For the blade parts, the compressor blade dimensions are within the print envelope constraints and the material is available and compatible with the Arcam Q20 printer. However, while the combustor blade dimensions are within the print envelope constraints, the specified material is incompatible with the printer.

However, while the combustor blade dimensions are within the print envelope constraints, the specified material is incompatible with the printer. For the fuel nozzle, it is also within dimensional constraints and made with a material that is available and known to be compatible with the printer.Figure 7: Feasbility Analysis of Aircraft Parts for 3D PrintingThe post processing capability, denoted K in the framework, is an important criterion in the F(p,T). Post processing involves heat treatment to yield the strength and machining to achieve surface finish and tolerance requirements of the printed part. Without the required post processing capability, the printed part will invariably not meet specifications even with technology ready 3D printers and materials. For example, in the case of the compressor vane blade, the feasibility of 3D printing may be greatly affected if the design of the blade contains holes for film cooling.

For example, in the case of the compressor vane blade, the feasibility of 3D printing may be greatly affected if the design of the blade contains holes for film cooling. Since the blade usually requires high surface finish, post processing the surface is almost always necessary. This requires sufficient machining capability, such as in-depth knowledge of electrochemical machining, so as not to block the holes or corrode the part though the process, rendering the part unusable. This is one of many examples of why post processing capability is germane in the feasibility analysis for 3D printing. Finally, parts that satisfies the F(p,T) requirements will pass the feasibility filter and undergo part value analysis.19705.4. Value Analysis of 3D Printed PartThe parts will be evaluated for part level value, V(p,T), based on the duration and cost to m).

Value Analysis of 3D Printed PartThe parts will be evaluated for part level value, V(p,T), based on the duration and cost to m). These metrics are manufacture for prototype intent (dp compared with current known duration and cost figures using existing manufacturing practices. A part that demonstrates superior duration and cost with 3D printing would be a good candidate for technology insertion. Conversely a part that lacks duration or cost advantages when 3D printed e) often carries should not be eliminated automatically. For a prototype intent part, the duration (dp e). The part level value, V(p,T), for a production intent part may be more weight than the cost (cp m) may hold similar importance.

For a prototype intent part, the duration (dp e). The part level value, V(p,T), for a production intent part may be more weight than the cost (cp m) may hold similar importance. However, even in the m) and cost (cp more complex as duration (dp scenario that a part is costlier to make, it may still deliver systemic benefits that are often overlooked in the local part level value analysis.e, cpe) and production intent (dpm, cpThe systematic value analysis, V(S), aims to uncover the systemic benefits of duration (ds) and cost (cs) for the assembled component. This analysis accounts the total time and cost to produce the assembled component with the 3D printed part and benchmarks those metrics against those of the current produced component.

This analysis accounts the total time and cost to produce the assembled component with the 3D printed part and benchmarks those metrics against those of the current produced component. A component that reveals a duration (ds) or cost (cs) saving may generate a discussion if benefits warrants a part to be 3D printed.An example of systematic value would be the fuel nozzle where its complex geometry has traditionally required multiple investment casted parts that are later welded to join the parts. The final part is both cost and labor intensive. When 3D printed, the part can be produced in a single run and without the need of welding, delivering a stronger part while requiring less labor. Identifying similar sub-assembly parts would potentially uncover tremendous cost and duration savings for the stakeholders.Expected Value of Framework Implementation6.1.

Identifying similar sub-assembly parts would potentially uncover tremendous cost and duration savings for the stakeholders.Expected Value of Framework Implementation6.1. Complement to Value AnalysisSeveral techniques have been in regular use to account for and calculate the potential value (e.g. material cost reduction and supply chain robustness) upon insertion of 3D printing. These techniques can be used to analyze potential benefit, yet in some cases do not address why hidden assumptions or ill-structured costs might hinder organizations from adopting the new processes.[28]–[30] As a complement, this paper’s framework emphasizes the linked consideration of background knowledge of processes along with the foreground knowledge in product design and engineering. As discussed earlier (Suzuki et.

As discussed earlier (Suzuki et. al.), background knowledge is the tacit knowledge embedded in an organization’s experience while foreground knowledge is the explicit knowledge shown in part design and production, such as engineering drawings and specifications.This framework complements value analysis by proposing a proactive approach to stimulate and address these background knowledge areas as part of exploration of value analysis for a portfolio of parts. Many past assumptions and experiences with traditional manufacturing may inform and differ from the direct value analysis; these concerns in the organization need to be addressed to allow initial steps and early pilots of the new process.19716.2. Readiness AssessmentThe framework also generates insights regarding the technology readiness and firm readiness to adopt 3D printing for the prototype and manufacture of a part.

Readiness AssessmentThe framework also generates insights regarding the technology readiness and firm readiness to adopt 3D printing for the prototype and manufacture of a part. The feasibility analysis of the part (P) in combination with the 3DP Technology (T) shows technology readiness of producing a particular part. This feasibility analysis considers the likelihood of satisfying the part attributes with a specific 3D printing process and material availability and can be viewed as an another example of foreground knowledge. However good technology readiness does not necessarily translate to high feasibility of producing a part. The tacit knowledge that is embedded in manufacturing routinely dictates the cost and quality of the finished article. An example of this tacit knowledge is the ability to identity materials and apply proper heat treatment to result the required yield strength.

An example of this tacit knowledge is the ability to identity materials and apply proper heat treatment to result the required yield strength. This tacit knowledge is often a core know-how and competitive advantage for a firm and thus considered the firm readiness. The framework considers this readiness (K) in the feasibility analysis, allowing the stakeholders differentiate and focus on the different readiness factors.6.3. Reverse Analysis of Value to PartsThe framework can potentially produce reverse analysis of value to determine the parts. This approach begins by establishing a desired cost and duration improvements, and the framework will identify the parts that have various feasibility of accomplishing the targets. To enable reverse analysis, the feasibility analysis F(p,T) should first be completed, along with the local value analysis V(p,T) and systemic value analysis V(S).

To enable reverse analysis, the feasibility analysis F(p,T) should first be completed, along with the local value analysis V(p,T) and systemic value analysis V(S). The reverse analysis introduces a top-down approach to value analysis, starting from the cost and duration targets, giving the different perspective that some stakeholders may require.ConclusionIn building this framework and in interviews with experienced manufacturers, several keyinsights about this approach were gained.The part by part consideration of 3DP feasibility requires technical know-how in the design, prototyping, pre-processing, production, post-processing and quality characteristics of the part in consideration. Experienced manufacturers possess this know-how for their baseline traditional processes, yet consideration of part by part in a BOM would simply be a stepwise, tedious, and very slow search. The required tacit knowledge is typically spread across multiple teams and locations.

The required tacit knowledge is typically spread across multiple teams and locations. Therefore, accelerated adoption should not proceed part by part. Whereas early in our framing of the approach we assumed an algorithmic selection of 3DP part candidates, deeper consideration pushed us to recognize the limits of a purely automated approach.This method’s emphasis on a “human in the loop” exploration of 3DP insertion recognizes that tacit knowledge of baseline processes, embedded from years if not decades of organization experience, need be stimulated to expose relevance (or not) of the accepted wisdom of the firm to the new process. Engaging existing product and process experts in the field helps to generate explicit knowledge – in a form that can be shared, tested, and questioned across teams. This social process may be necessary for full consideration of 3DP which displaces proven processes at quality,1972cost, and schedule well understood from past experience.

This social process may be necessary for full consideration of 3DP which displaces proven processes at quality,1972cost, and schedule well understood from past experience. The awareness and willingness of human teams to expose and unlearn past assumptions was seen as a significant barrier to early steps of considering 3DP.Thus the framework stimulates a first step by filtering candidate parts and subsystems with up to date combination of the latest material and 3DP tools. Yet the authors now view the framework primary purpose to stimulate team of teams dialogue more quickly and efficiently. The active engagement of cross-functional experts in search of the trade space is meant to accelerate their situational awareness of the systemic feasibility and value, rather than partwise feasibility.7.1. Limits and future workThis paper presented a challenge and framed a method for improved adoption of 3DP in organizations with significant legacy know-how and a large part portfolio.

Limits and future workThis paper presented a challenge and framed a method for improved adoption of 3DP in organizations with significant legacy know-how and a large part portfolio. Organizations starting with “green field” development may be able to rapidly displace old or acquire new product development teams and know-how, thus avoiding the sociotechnical barriers that this method addresses.In this paper an example was described as a basis for thinking through the method. While more detailed algorithms for filtering and prioritizing parts were developed by the authors, these have not yet been validated and thus are not presented in this paper. Future work may include deployment of a prototype system based on the method with observation of the adoption rates across teams. Formal linkage to existing value analysis and value engineering methods would strengthen the cost and duration calculations in the portfolio filters and prioritization.7.2.

Formal linkage to existing value analysis and value engineering methods would strengthen the cost and duration calculations in the portfolio filters and prioritization.7.2. AcknowledgmentThe authors would like to thank our colleagues at the Massachusetts Institute of Technology (MIT), the University of Tokyo, and Mitsubishi Heavy Industries America (MHIA) for their support and feedback on this work.References[1] C. Lindemann, T. Reiher, U. Jahnke, and R. Koch, “Towards a sustainable and economic selection of partcandidates for additive manufacturing,” Rapid Prototyp. J., vol. 21, no. 2, pp. 216–227, 2015.[2] D. S. Thomas and S. W. Gilbert, Costs and Cost Effectiveness of Additive Manufacturing. NIST, NationalInstitute of Standards and Technology, US Department of Commerce, 2014.[3] A.

 Printing Thin Skinned Composites to Achieve the Strength-to-Weight Ratio of Aluminum. David E. Fly11 Department of Engineering and Technology, University of Wisconsin–Stout Menomonie,WI. USAABSTRACTKevlar and stainless steel mesh reinforcements were added using epoxy to 3D printed ABS-M30 thin skins, thereby making a composite structure with significantly improved mechanical properties over that of the 3D printed plastic alone. These additive manufactured composites have a strength to weight ratio that is comparable to solid aluminum. Flexural 3-point bend tests and Charpy Impact tests were conducted. Experiments were conducted that were designed to characterize the influence of adding Kevlar to the composite structure and also the influence of pre-mixing glass microspheres into the epoxy. These new additive manufactured (AM) composites are an attractive choice to designers attempting to reduce weight because any 3D printed shape can be reinforced in this manner.

These new additive manufactured (AM) composites are an attractive choice to designers attempting to reduce weight because any 3D printed shape can be reinforced in this manner. Additionally, actual production time is less than 3D printing a fully solid component.Keywords— Composites; strength-to-wieght ratio; additive manufacturing, 3D printing.INTRODUCTIONThe phrases 3D printing and additive mmanufacturing are often used interchangeably even though they have slightly different meanings. Technically speaking, 3D printing is only one type of additive manufacturing (AM) but since these technologies have gone mainstream in public use the phrase 3D printing has been used in a broad sense for all these technologies. Fused deposition modeling (FDM) is the extrusion of a thread of plastic deposited in shapes created from a sliced 3D model. Each shape becomes one layer and successive layers build a 3D plastic component.

Each shape becomes one layer and successive layers build a 3D plastic component. Most people would refer to FDM as the 3D printing of plastic even though there are many other technologies that 3D print plastic such as selective laser sintering and inject printing. FDM was used in this research and is used synonymously with 3D printing in this paper.The advantages of 3D printing plastics are extensive and well known to most readers of solid freeform fabrication symposium conference papers. Two key disadvantages are the speed of printing and the fact that plastics are typically not very strong relative to other materials. This research explores and idea that improves upon those two disadvantages.By 3D printing two thin skins that are designed to interlock, a hollow 3D structure can be printed in less time than if the full 3D structure were printed entirely as a whole. Additionally, by printing thin skins that interlock, reinforcements can be added internally using epoxy.

Additionally, by printing thin skins that interlock, reinforcements can be added internally using epoxy. These internal reinforcements can include fiberglass mesh, steel mesh, carbon fiber, aramid fabric (Kevlar), aluminum, titanium. These can be used for external reinforcements as well, and/or integrated between layers. A variety of adhesives can be used although epoxy is likely the mostcommon. The adhesive can contain fiber additives, be manually applied, or vacuum infused. Some authors refer to these technologies as hybrid composites. Figure 1 is a photo of two interlocking thin skins with stainless steel mesh inside, ready for adhesive.Figure 1 - Interlocking 3D printed thin skins with stainless steel mesh inside.Figure 2 - Interlocking 3D printed thin skins showing the opening on the right side for injectingthe adhesive by syringe.RESEARCH OBJECTIVES1. Develop methods for fabricating reinforced composites using 3D printed or FDM thin skins.2.

Develop methods for fabricating reinforced composites using 3D printed or FDM thin skins.2. Compare the strength to weight ratio of reinforced composites to that of solid aluminum.3. Conduct experiments using Charpy impact testing (to standards of ISO-179) that examine the influence of internal reinforcement and hollow glass microspheres in epoxy.4. Conduct flexural 3 point bend testing (to standards of ISO-178) on reinforced composites.EXPERIMENTSThree experiments were proposed:1. ANOVA using impact strength as the response variable to compare internal reinforcements of fiberglass mesh versus stainless steel mesh using hot glue.2. Blocked ANOVA using impact strength as the response variable to examine epoxy with glass microspheres as additive into the epoxy.3.

Blocked ANOVA using impact strength as the response variable to examine epoxy with glass microspheres as additive into the epoxy.3. ANOVA using flexural strength as the response variable to examine epoxy with various reinforcement types of stainless steel mesh, Kevlar fabric, carbon fiber.Figure 3 - Specimens being prepared by injecting the adhesive in the first experiment.RESULTS AND DISCUSSIONResults of the first experiment showed potential, an ANOVA using impact strength as theresponse variable to compare internal reinforcements of fiberglass mesh versus stainless steel mesh using hot glue. The analysis of variance did yield evidence of significant differences in the reinforcement types, with p-values even lower than 0.0005. This full factorial experiment showed the strongest composite construction to be the combination of fiberglass mesh and stainless steel mesh.

This full factorial experiment showed the strongest composite construction to be the combination of fiberglass mesh and stainless steel mesh. Figure 4 is a box plot of the results of the first experiment.1.81.61.41.21.00.80.60.40.2Fiberglass_1Stainless101010Energy Consumed (ft*lb)Energy Consumed (Hot Glue)Figure 4 - Box plot of the results of the first experiment.Variation in these data for each factor level combination was noticeable. It is assumed that much of this variation was the influenced by the presence of internal voids in the adhesive. These voids formed during fabrication in part due to the viscosity of the adhesive and the rate at which the adhesive filled the internal cavity.

These voids formed during fabrication in part due to the viscosity of the adhesive and the rate at which the adhesive filled the internal cavity. These internal voids are visible in Figure 5, a photo of the specimens fabricated for the first experiment, prior to the impact tests.Figure 5 – Specimens fabricated for the first experiment, inconsistent filling of the internal cavityis visible in all factor level combinations.Results of the first experiment were also compared to similar measurements of the impact strength of 6061-T6 aluminum. The first experiment was conducted to the international standards of ISO-179-1. This is a pendulum impact test that measures the amount of energy consumed by the specimen as a result of the pendulum impacting the specimen. A larger amount of energy consumed indicates a stronger specimen. Aluminum specimens were prepared to the same dimensions as that of the 3D printed specimens (3.1mm x 15mm x 77.5mm). The span of the supports was 62mm.

Aluminum specimens were prepared to the same dimensions as that of the 3D printed specimens (3.1mm x 15mm x 77.5mm). The span of the supports was 62mm. Dividing the impact energy consumed by the weight of the specimen yields a strength to weight ratio. Calculations are as follows:S/W ratio of a Stainless Steel Mesh with Fiberglass Mesh : 1.75 J / 54.9 Newtons = 0.0318 J/NewtonS/W ratio of 6061-T6 Aluminum :5.317 J / 88.80 Newtons = 0.0599 J/NewtonOne conclusion of the first experiment were that reinforcing 3D printed thin skins has potential and attention needs to be given to the way the adhesive is inserted into the cavity in order to minimize the internal voids. The second conclusion of the first experiment was that 3D printed thin skins reinforced internally might be able to achieve the strength to weight ratio of a strong aluminum. This will depend on the adhesive type, consistency, and reinforcement type.

This will depend on the adhesive type, consistency, and reinforcement type. It is noteworthy that carbon fiber was not used in the first experiment and it has good potential for a lightweight material with high strength. Carbon fiber was used in the third experiment of this research.The second experiment of this research was a blocked ANOVA using impact strength as the response variable to examine epoxy with hollow glass microspheres as additive into the epoxy. The two 3D printed materials used were ABS-M30 and Ultem. The hollow glass microspheres (glass bubbles) were intended to reduce weight without a loss in strength. Unfortunately, the internal voids formed during the first experiment were even more pronounced in the second experiment. These voids prevented meaningful data collection as the specimens would always break at one of the large voids. The voids were so inconsistent that often the impact reading from a test was not even on the scale of the instrument.

The voids were so inconsistent that often the impact reading from a test was not even on the scale of the instrument. A decision was made to stop the data collection during the second experiment since it was of no real value.Viscosity of the epoxy was a factor, and adding the hollow glass microspheres produced anoticeable change to the viscosity of the epoxy. Figure 6 is a photo of the specimens for the second experiment. Figure 7 is an image that reveals the size distribution of the glass microspheres.Figure 6 – Specimens fabricated for the second experiment, internal voids are visible.Figure 7 – Image shows the size distribution of the glass microspheres.The third experiment of this research was an ANOVA using flexural strength (as per ISO- 178) to examine the influence of three reinforcement types when fixed in position with epoxy only. The three reinforcement types were stainless steel mesh, aramid fabric (Kevlar), and carbon fiber.

The three reinforcement types were stainless steel mesh, aramid fabric (Kevlar), and carbon fiber. As with the previous experiments, internal voids were visible in the specimens and the data indicated a corresponding variation in the data. In this third experiment, full data was collected but the variation was too large to make meaningful conclusions from these results. These inconclusive data are summarized in Figure 8.Figure 8 – Results of the third experiment produced inconclusive comparisons.SUMMARY AND CONCLUSION3D printing thin skins to reinforce them as hybrid composites was shown to be feasible. Preliminary data have achieved approximately half the strength to weight ratio of 6061-T6 aluminum and those specimens did not use carbon fiber. Three experiments were planned and two of those yielded data. Internal voids in the adhesive were an ongoing problem that produced unacceptably large variations in data measurements for both impact and flexural strength.

Internal voids in the adhesive were an ongoing problem that produced unacceptably large variations in data measurements for both impact and flexural strength. Visible differences in the viscosity of the adhesive seem to strongly influence the formation of internal voids.Future research has begun which will 3D print thin skins with louvers. A thickened epoxy will be used such that during fabrication, the adhesive over fills the thin skins in a way that allows the adhesive to extrude outward through the louvers. This shows promise for the ability to fabricate these hybrid composites with minimal internal void formation.ACKNOWLEDGEMENTThe author would like to acknowledge his employer, the University of Wisconsin Stout.REFERENCES[1] Espalin D., K. Arcaute, E. Anchondo, A. Adame, F. Medina, R. Wicker, T. Hoppe, R.Wicker. Analysis of bonding methods for FDM-manufactured parts. Solid Freeform Fabrication Symposium. 2010.[2] Jana S., B.R. Hinderliter, W.H.

 PRINTING WITH NATURAL FIBER REINFORCED FILAMENTJ. I. Montalvo N*, M.A, Hidalgo*Department of Energetics and Mechanics, Universidad Autonoma de Occidente, Cali, ColombiaAbstractAn initial study of 3d printing with compound filament using different plastic matrices and sugar cane bagasse as the filler was conducted. In order to do this, a reverse engineering process was made to several 3d printer extruders to determine how to change the extruder in order to be able to print with the filament. To obtain the filament, a plastic extruder was modified to obtain a compound filament of 1.75 mm using a 3x4 design of experiments with the factors percentage of fiber (10% 20% 30%) and type of matrix(PE,PP,ABS,PLA).

The filaments obtained were tested to determine the mechanical properties and finally were used in a 3d printing to compare results.IntroductionAdditive manufacturing technologies have the advantage over traditional technologies to produce fast, flexible and competitive parts, prototypes, molds, dies, or custom end products directly from digital information[1]. Additionally, they have the advantage of reducing product development times, and most importantly fabricate designs and features unmatched by other methods of manufacturing[2].The downside of the pieces obtained by rapid prototyping technologies are that the products obtained do not have the same mechanical properties than those obtained by traditional manufacturing processes. Additionally, the filament used in the process is not locally produced or distributed in Colombia, making the prices of the process higher for massive use.

Additionally, the filament used in the process is not locally produced or distributed in Colombia, making the prices of the process higher for massive use. The proposal is to study the fabrication of a natural fiber reinforced composite of thermoplastic matrix reinforced with natural fibers, a combination that has shown in other applications to have higher mechanical performance than when using only the thermoplastic material [3], and investigate their application in the field of rapid manufacturing, specifically for Fused Deposition Modeling, FDM. To this, a design of experiments for the material was made, the material was mechanically characterized, and a reverse engineering process of some additive manufacturing equipment was made.Materials and MethodsReverse EngineeringThe starting point for this research was the process of reverse engineering to a FDM 3d printing machine. The machine chosen for this process was a 3d Thing-o-Matic printer.

The machine chosen for this process was a 3d Thing-o-Matic printer. This was a pioneer in the field of low-cost 3D printers from the Makerbot company, whose dimensions are approximately 30x30x40 cm, with a build area of approximately 0x10x10. This printer was one ofthe first fully automated commercial 3D printers, by software all printing parameters can be controlled, and once the print begins, no human intervention is required until the process ends. This makes it a machine suitable for the process of reverse engineering, well it is very complete, and additionally low cost.Reverse engineering is a systematic process of dismantle where it is intended to list the problems and design opportunities, and to disassemble, measure and analyze the modules or operational subsets, to extract a list of primary functions, and components related to each functions, to understand operation.

Figure 1 shows some of the functional subassemblies of the machine.a)b)c)d)e)f)a) Control System b)Hot Platform c)Extruder d)X axis Mechanism e) Y axis Mechanism and f)Z axis mechanism.Figure 1. Functional Subsets of the Commercial Thing o Matic printer.After the reverse engineering process, functional subsets were identified. In order to generalize the common functions for 3D printers, a similar process was also realized with other 3Das shown in Figure 2, and although the reverse engineering process was not applied completely, it served as a reference point for identifying these functional subsets. The printers chosen range in price, quality and brand, such as Stratasys Dimension 1200es, which is a professional industrial printer whose market value is around thirty thousand dollars, and the Mendel Prusa, an amateur kit for hobbyists whose cost is around six hundred dollars.

Other printers are in the medium range, about two thousand dollars each, but widely use in the growing industry of personal 3D printing.After realizing the reverse engineering process, and experimenting with the referenced machines, the following subsets where identified: structure or framework, positioners in each of the axes X, Y and Z, one printing platform extrusion mechanism and a control system.a) Makerbot Replicator 2 b) Makerbot Replicator 2x c) Makergear d) Prusa Mendel e) Stratasys Dimension 1200es Figure 2. Reference 3d Printers.Generally, the majority of printers have a steel structures with calibrated axes used as guidelines for the displacement in each axis of motion. Additionally, pulleys and belts are used to transmit movements in the x and y plane, and worm gears are used for lifting the platform in the z axis.

Additionally, pulleys and belts are used to transmit movements in the x and y plane, and worm gears are used for lifting the platform in the z axis. Extrusion mechanisms vary from one machine to another, as some are single headed while others are double, some have the motor directly coupled to the toothed pulley feeding the filament, while others have a gear system, and operating temperatures for the head range between 25 and 300 °C. The platforms generally have temperature control for printing with ABS, and the materials of these platforms can be plastic, metal or glass. Table 1 list the functional subsets identified in the machines studied.StructureMovement MechanismExtruderControlStructure Filament Support Hot Plate X Axis Y Axis Z Axis Filament Speed Temperature Control Electronics Motors Firmware and SoftwareTable 1.

Functional Subsets in 3d printers.Since the main goal of this research was to use a NFRC in FDM printing, the extruder head or printing block of most of these printers was also studied using the reverse engineering process, to clearly identify its functions and components and be able to learn based on the knowledge gained from existing machines. In Figure 3 the extruder blocks used in the study are shown.a)b)c)A. Thingomatic. B. Stratasys. C. Makergear.Figure 3. Reference Extruder Blocks.After realizing the reverse engineering process to each of the heads in order to obtain a better understanding of them, the functional decomposition begins. Initially the head is visualized a whole, both inputs and outputs for the system are determined, and summarized in a black box diagram. Then the functional decomposition show the relationship between functions in order to transform the input into the output.

Design of ExperimentsIn order to obtain the 1.75mm in diameter filament to be used in a 3d printer, an extruder must be used. An old extruder used to make plastic hose was modified and reconditioned in order to obtain the filaments. After the conditioning of the machine was done, the composite fabrication process started. The polymeric matrixes were obtained in pellets, and the natural fiber used was obtained pulverized with the help of SENA, Astin Center. The filament was manufactured in the plastic transformation Lab of the GITEM investigation group at Universidad Autonoma de Occidente. Figure 6 shows the process of obtaining the filament. The proper quantities of matrix and filament where weighted and manually mixed, to be feed into a HOT PLATE PRESS where a standardized cycle was used in order to apply controlled temperature and pressure to homogenize the composite.

The output of this process was a rigid board which was then pelletized in a mill, and then feed to the extruder in order to extrude the filament.Figure 6. Filament fabrication process.This process turned out to be the most important for the design of experiments because of the experimental validation that in order to obtain a good filament, other additives and chemical treatments must be used for a successful extrusion. Additionally, to obtain an homogenous material, a twin screw extruder must be used, and thanks to the conclusions of this project, it was requested and will be used in future investigations.Table 2 shows the summary of the design of experiments for the extrusion process.

Extrusion property refers to the ability to get the filament out of the nozzle of the extruder, speed refers to the speed with which the material is allowed to pull to reduce its diameter after leaving the nozzle, and fragility relates how fragile is the filament obtained.Fiber Percentage %Matrix Property Extrusion Speed Fragility Extrusion Speed Fragility Extrusion Speed Fragility Extrusion Speed FragilityPEPPPLAABS0% 10%20%30%Yes Yes Yes Yes NO Medium Low NO - Low Low Yes Yes Yes Yes High Medium Low High Low Low Low Yes Yes Yes Medium Low No High Yes Medium - - HighMedium Yes No - No - -High - No No-Table 2. Extrusion Results.When trying to extrude the ABS, very brittle filament was obtained. The literature review revealed that commercial ABS filaments ranges between 75% and 90% ABS, and the remainder is a co-polymer of styrene and acrylonitrile (SAN) and dyes or other additives which increase the ductility which reduces fragility[4][5].

Similarly, for the PLA it was also found that the percentage composition ranges from 85% to 90% polymer and the remainder are additives which improve its properties[6][7]. PP and PE are not widely used for 3D printing as they are very flexible and its adhesion properties are not very good [8]. Moreover, their shrinkage coefficients are much greater than those of ABS and the PLA [9], which produce deformed parts with design measurements deferring from final parts. However, these materials are easily extruded, and it can be observed that when increasing the percentage of fibers, extrusion speed must be slowed in order to obtain filament.

However, these materials are easily extruded, and it can be observed that when increasing the percentage of fibers, extrusion speed must be slowed in order to obtain filament. The literature research also explains this as increasing fiber content increases the brittleness of the composite [10] and decreases if ductility.Afterwards, it was observed that the mixture of PP with 20% sugarcane bagasse was the composite which best behaved at the time of extrusion, and in agreement with the ASTIN- SENA, and the Laboratory of Physical Tests polymers, physical and mechanical tests were performed to determine their behavior. Annex B shows the results of the tests conducted, five samples of each compound of polypropylene, which further were chemically treated to try to better adhesion of the fibers to the matrix.

Annex B shows the results of the tests conducted, five samples of each compound of polypropylene, which further were chemically treated to try to better adhesion of the fibers to the matrix. Figure 7 shows a summary of the test, where it can be observed that the fibers are working as filler and not as reinforcing agents well pure polypropylene properties are better than the compound.A) 3 point Flexion PP B) PP Tension. C) 3 point Flexion PP-20% Bagasse D) Tension PP- 20% Bagasse Figure 2. Mechanical Flexion and Tension Results.Manufacturers and sellers of fiber-reinforced composites recommend not to print with nozzles less than 0.5mm diameter [11][12][13]. An experimental setup where the filament obtained was printed in air until a steady flow without changes in the output was obtained was performed. It was tested with nozzles of 0.4mm, which is the standard nozzle used for printing with polymeric material, and was gradually increasing by 0.05mm to 0.8mm in diameter.

It was tested with nozzles of 0.4mm, which is the standard nozzle used for printing with polymeric material, and was gradually increasing by 0.05mm to 0.8mm in diameter. Additionally, a line pattern was printed to check resolution and bonding between the layers, varying the diameter of output parameter in the software to compensate for changes in the nozzle.Table 3. Nozzle TestsNozzle Ø (mm) 0.40 0.45 0.50 0.55 0.600.650.70 0.75Continuous FlowNo No No Yes YesYesYes YesTime without clogs (s) 3 4 5 6 101010 10ObservationsVery Thin, some blank spaces Some clogs, separated lines Well-formed lines and separations. Constant flow, overlapping lines. Overlapping lines, Material flow with motor off.0.80Yes10Table 3 shows a summary of tests performed to determine the best nozzle diameter depending of the composite behavior. For diameters smaller than 0.5mm, the nozzle tends to become clogged generating an irregular flow, and some spaces remain unprinted.

For diameters smaller than 0.5mm, the nozzle tends to become clogged generating an irregular flow, and some spaces remain unprinted. For diameters between 0.5mm and 0.7mm, and the flow is more constant with almost no obstructions and the printed lines are well defined, but for diameters larger than 0.7mm, the flow is abundant and lines overlap each other, and it was observed that the material was flowing out even if the motor was turned off, only due to gravity, which is not suitable for the printing process where it is required to control the times when printing is required. Because of these considerations, the diameter of the nozzle for the printer extruder head was 0.6 mm, 50% larger that of the nozzles used to print polymer material without fibers. Figure shows the filament before and after extrusion by the printer.Figure 8. 1.75 mm filament (top) and extruded filament (bottom)Figure 9 shows a printed sample using the filament, at 1x, 10x and 20x.

1.75 mm filament (top) and extruded filament (bottom)Figure 9 shows a printed sample using the filament, at 1x, 10x and 20x. The filament layers and line overlapping can be observed. The presence of natural fibers can be clearly observed, and it is demonstrated that the filament can be used to obtain a 3d part since the material layers are joined and the material is deposited in the necessary points in order to obtain the modeled part. Additionally it can be seen that the size of the fibers in the material is not constant, explaining why the gear drive and tensioning system must be used.Figure 9. 3d printed sample. (Top). 10x close up (Bottom Left) and 20X (Lower Right)The material chosen for the thermoplastic matrix is very flexible, so the parts that have been printed are 2D with volume, but further investigations are in process to obtain a functional filament using PLA and ABS.

In order to obtain this, other additives are needed in the extrusion process as explained above, and additionally a double screw extruder with volumetric dosing is suggested. However, these developments are very promising, because it verifies that the head is able to print composite materials manufactured locally, providing autonomy and opening a research line to obtain new products or prototypes of composite materials using additive manufacturing.CONCLUSIONSThe presented work shows the design and development process followed in order to obtain a natural fiber reinforced composite filament suitable for using in 3d printing by FDM. The reverse engineering process was a very useful tool in order to understand the process, functions, and parts that make up a 3d printer and how several recommendations must be followed in order to print with natural fiber reinforced composites.

lding and MillingDirect Prototyping of MetallicYong-Ak Song*, Sehyung Park*, Kyunghyun Hwang**, Doosun Choi**, Haeseong Jee***Korea Institute ofScience and Technology KIST, CAD/CAM Research Center ** Korea Institute ofMachinery and Metals KIMM *** Hongik University, Dept. ofMechanical EngineeringAbstract:Welding has been used for the direct fabrication of metallic prototypes and prototype tools by several research institutes. Since welding alone is not able to deliver the accuracy and the surface quality needed for prototype tools, especially for injection molds, a combination with conventional machining is necessary. In this paper, welding and 5-axis milling are combined together for the direct fabrication of metallic parts. For welding, conventional CO2 arc welding is used. Test parts with conformal cooling channels an~ undercuts demonstrate the technological potential of this process combination for rapid tooling applications.1.

Test parts with conformal cooling channels an~ undercuts demonstrate the technological potential of this process combination for rapid tooling applications.1. IntroductionWith an increasing demand to the prototype tools in regard to dimensional accuracy as well as to mechanical strength, several rapid tooling techniques have been developed and applied in issue of industrial application [1, 2]. Although they seem quite promising, dimensional accuracy remains for the currently used rapid tooling techniques. Usually, each extra step required for the conversion process towards the final part implies a deterioration of it would be best if the prototype tool can be its dimensional accuracy. As matter of fact, manufactured directly without any secondary process.the criticalFor this purpose, several approaches for the direct prototyping of metallic tools are being developed.

As matter of fact, manufactured directly without any secondary process.the criticalFor this purpose, several approaches for the direct prototyping of metallic tools are being developed. The common feature of the processes consists in direct melting of metal with a focussed laser beam, an arc or a plasma. The material used is normally supplied either in form of powder or wire. Because of the complete melting, however, the obtained accuracy as well as the surface quality of parts are generally lower than that of machined parts.To overcome this difficulty, the process combination with a substractive technique such as conventional milling can be an appropriate solution. This kind of process combination has been already successfully implemented in shape deposition manufacturing [3,4] and laser generating [5]. In this investigation, it is aimed to discuss some fundamental technological aspects of the process combination for its further development.4952.

In this investigation, it is aimed to discuss some fundamental technological aspects of the process combination for its further development.4952. Experimental ProcedureThe test facility consists of a 5-axis vertical milling machine and an arc welding equipment, Fig. 1. The welding gun attached to the spindle can be moved in z direction. The substrate plate is fixed on a rotary table mounted on the linear x-y table of the machine tool.5-axis machine.-r====J.;~ substrate platerotary tableFig. 1: View of the test facilityAs shown on the right side of Fig. 1, the substrate plate is placed directly beneath the welding gun at first, and then moved to the milling tool after welding operation is completed. In milling operation, the top as well as the sides of the bead are machined to a specified height and width before depositing a new layer upon the previous one.

In milling operation, the top as well as the sides of the bead are machined to a specified height and width before depositing a new layer upon the previous one. This procedure repeats until part is completed.Among the significant process parameters of 3D welding and milling are voltage and current as well as travelling speed of the substrate plate. Prior to the experiment, the relative position between the welding gun and the substrate plate and the flow rate of shielding gas were set according to the result of the previous investigation and kept constant during the experiment. For welding operation, a carbon steel wire (AWS A5.l8) with a diameter of 0.9 mm was used. The substrate plate was also made of a carbon steel.496As an alternative to arc welding, laser beam welding can be used in the experiment. Since laser beam can be more precisely focussed than welding arc, more accurate parts can be built.

Since laser beam can be more precisely focussed than welding arc, more accurate parts can be built. In addition, focussed laser beam allows a less heat input into the part, thus making the part less susceptible to distortion. However, the deposition rate of laser beam welding is generally lower than that of arc welding. Besides, the utilization of laser requires a more complicated beam delivery mechanism and a more precise wire feeding system. Considering these features, laser beam welding is more suitable for manufacturing small and precise parts rather than large and bulky parts. A selection between arc and laser beam welding should be made depending on the geometrical complexity and the size of part.3. Results and Discussion3.1 Single layer experimentStarting from the single bead experiment, the influence of voltage, current and travelling speed on the formation of welding beads was investigated. Voltage and current were varied in ranges of l6V-36V and 60A-140A, respectively.

Voltage and current were varied in ranges of l6V-36V and 60A-140A, respectively. Optimizing the parameters in the ranges above at a constant travelling speed allowed to deposit single connected beads instead of loose spheres. It was shown that the bead size heavily depends on the travelling speed of the substrate plate, Fig. 2. When the travelling speed exceeds vs= 600 mmls, however, the height decreases only slightly down to 0.5 mm, whereas the width linearly decreases to 1.8 mm.5o-width40- heightE 3 .s .c:....."0 .~2..,..; .c: 0)'w .c:1voltage 24V - current 120A0.20.40.60.81.01.21.4travelling speed [100mm/s]Fig. 2: Height and width of single beads at different travelling speeds497The result in Fig. 2 also shows that varying the travelling speed between Vs = 200 mm/s and 1200 mm/s constitutes an effective method to control the size of beads.

2 also shows that varying the travelling speed between Vs = 200 mm/s and 1200 mm/s constitutes an effective method to control the size of beads. Compared to the other process parameters such as voltage and current, the travelling speed is relatively easy to change during the process. For this reason, if the bead size is required to be changed during the process, the travelling speed should be varied.3.2 Multiple layer experimentThe fundamental difference between a single layer and a multiple layer is the condition of heat conduction from the processing area. In case of the single layer, the heat can dissipate relatively fast to the substrate plate, whereas in case of the multiple layer, the heat is only slowly conducted to the substrate plate. This is partly due to the reduced cross-sectional area and partly due to the dissipated heat from the previous layers.

This is partly due to the reduced cross-sectional area and partly due to the dissipated heat from the previous layers. Consequently, the heat is constantly accumulated in the part until an equilibrium between heat input and dissipation is achieved. Since this accumulation of heat influences the welding result, variation of the process parameters is necessary depending on the height. An effective method to obtain a constant result would be adapting the travelling speed according to the actual height of building part.In Fig. 3, thin walls made of 8 layers are shown without and with machining. In case of the part without machining between depositing layers, the instability of arc welding process caused a defect in the middle of the beads. Initially small in size, the defect in the second layer influenced the height of the next deposited bead and, thereby, amplified itself until a further deposition was no longer possible.

Initially small in size, the defect in the second layer influenced the height of the next deposited bead and, thereby, amplified itself until a further deposition was no longer possible. Therefore, machining the bead perpendicular to z direction is not only necessary to increase the part accuracy, but also to make the process more stable for the further deposition. When milling the deposited beads perpendicular to z direction, a change of the height was observed. The first four layers had a height of around 2.8mm, while the succeeding layers showed a decreased height of around Imm. This result is to explain with the changing cooling rate in the part, as mentioned in the previous paragraph.The analysis of the cross-section in the wall reveals a different microstructure depending on the height. The microstructure taken from the upper area of the wall shows enlarged grains with a spherical shape (see a in Fig. 3).

The microstructure taken from the upper area of the wall shows enlarged grains with a spherical shape (see a in Fig. 3). In each grains, a second phase can be observed which is only viable when the cooling rate is relatively low. Compared to that, the microstructure ofthe lower area contains relatively small grains as well as some dendrites (see b in Fig. 3). Such a microstructure results from rapid cooling due to the large amount of heat conduction to the substrate plate. In conclusion, the formation of microstructure also proves the changing cooling rate with increasing height. Therefore, an adjustment of the process parameters according to the height is necessary to maintain a homogeneous microstructure in the part. Otherwise, the parthas to be heat-treated.498without millingwith milling0: welding defectparameters: 24V, 120A, 200mm/ssize: 40mm x 1.5 mm x 22 mm10 umFig. 3: Thin-walled parts without milling and with millingIn Fig.

3: Thin-walled parts without milling and with millingIn Fig. 4, two inclined walls built with an offset value of 0.5 mm and 1 mm, respectively, are shown. After depositing 15 layers, the walls were machined with a tilted milling tool in one operation.+- before millingparameters: 24V, 120A, 200mm/soffset value: 0.5 mm, 1 mmlayer number: 15Fig. 4: Inclined walls with angles of 24° and 44°499Besides thin walled parts, 3D welding and milling also enables manufacturing solid parts with good surface and dimensional accuracy. In Fig. 5, a solid part fabricated with 3D welding and milling is shown. The shifting distance between each beads was 3.7mm. The test part contains an internal channel around the rectangular cavity. Theoretically, it would also have been possible to machine the internal channel of the part from a block and close the channel with welding afterwards. The cavity in the middle of the part could also have been completely machined after finishing deposition.

The cavity in the middle of the part could also have been completely machined after finishing deposition. However, it would have taken much more time because of the large volume to be removed. If the volume to remove were less than the deposited one (e.g., the walls of the internal channel were thicker than the channel itself), then machining would have been preferable to welding.This example implies how flexible the combination of welding as additive and milling as substractive technique can be applied for prototyping. To use the process effectively, however, an intelligent process planning is necessary taking the complexity and the size of features into consideration as well as the required mechanical properties. Based on this consideration, it has to be decided which features of the part to build with additive or substractive method.parameters: 24V, 120A, 200mm/s size: 70mm x 65 mm x 20 mmFig. 5: Solid test part5004.

5: Solid test part5004. ConclusionsAccording to the current state of the art, dimensional accuracy as well as surface quality of rapid prototyped metallic tools are still far behind that of conventionally machined tools. To increase the quality of rapid prototyped tools, a hybrid approach comprised of an additive and a substractive technique called 3D welding and milling is being developed.The combination with milling operation not only increases both the surface quality and the dimensional accuracy, but it also gives a great manufacturing flexibility. To use the process combination effectively, however, an intelligent process planning depending on the part geometry and its size is a prerequisite. In addition, a process control is required as the heat dissipation changes with increasing part height.5. References[1] S.

In addition, a process control is required as the heat dissipation changes with increasing part height.5. References[1] S. Ashley, "From CAD art to rapid metal tools", Mechanical Engineering, ASME, March1997, pp.82-87[2] Rapid Prototyping Report, CAD/CAM Publishing, June 1998 [3] R. Merz, F.B. Prinz, K. Ramaswami, M. Terk, L. Weiss,"Shape Deposition Manufacturing", Proceedings of the Solid Freeform Fabrication Symposium, University of Texas at Austin, August 8-10, 1994[4] J. R. Fessler, R. Merz, A. H Nickel, F.B. Prinz, "Laser Deposition of Metals for Shape Deposition Manufacturing", Proceedings of the Solid Freeform Fabrication Symposium, University of Texas at Austin, August 1996[5] F. Klocke, T. Celiker, Y.-A. Song, "Rapid Metal Tooling", Proceedings of the 4th European Conference on Rapid Prototyping and Manufacturing, Belgirate, Italy, June 13-15, 1995501502.

-inkjet Printing of Flexible and Stretchable ElectronicsJ. Vaithilingam*, E. Saleh, C. Tuck, R. Wildman, I. Ashcroft, R. Hague, P. DickensAdditive Manufacturing and 3D Printing Research Group, Faculty of Engineering, University of Nottingham, Nottingham, UK, NG7 2RD.AbstractInkjet printing of conductive tracks on flexible and stretchable materials have gained considerable interest in recent years. Conductive inks including inks with silver nano- particles, carbon based inks, inks containing poly (3,4-ethylenedioxythiophene) (PEDOT) doped with polystyrene sulfonic acid (PSS) are being researched widely to obtain a printed electronic patterns. In this study, we present drop-on-demand inkjet printing of conductive silver and PEDOT:PSS on a flexible and stretchable substrate. Process conditions for the inkjet printing of silver nano-particles and PEDOT:PSS were optimised and simple geometrical patterns (straight line and sinewave tracks) were printed.

Process conditions for the inkjet printing of silver nano-particles and PEDOT:PSS were optimised and simple geometrical patterns (straight line and sinewave tracks) were printed. Surface profile, surface morphology and electrical resistance of the printed patterns were examined. The printed silver patterns were observed to be highly conductive; however when stretched, the patterns did not conduct due to the origination of cracks. The measured conductivity for the PEDOT:PSS patterns was significantly lower than the silver patterns; however, they remained conductive when stretched for up to 3 mm. When flexed, PEDOT:PSS remained conductive for a lower radius of curvature (10 mm) than the silver. Among the printed patterns, the sinewave pattern was observed to be superior for flexible electronics application.1. IntroductionFabrication of flexible electronics is gaining significant interest and the market is rapidly growing.

IntroductionFabrication of flexible electronics is gaining significant interest and the market is rapidly growing. Printed electronics has the potential to eliminate the lengthy, material- intensive and expensive processes used traditionally [1]. Printing technologies including screen printing, soft lithography, flexography, gravure and inkjet are used for the fabrication of flexible electronics. Currently, flexible electronics are fabricated either by assembling a completed circuit on a flexible substrate or by fabricating the circuit directly on a flexible substrate [2]. Assembling electrical components to a substrate being a multi-step process, fabrication of flexible electronics all in a single-step would be beneficial.

Assembling electrical components to a substrate being a multi-step process, fabrication of flexible electronics all in a single-step would be beneficial. Inkjet printing of conductive materials over a substrate offers significant advantages including reduced wastage and ease of functionality [3].Inorganic metal based inks containing gold, copper, aluminium and silver are used for the fabrication of electronic devices including the RFID tags, transistor electrode and other passive components [4], [5]. However, due to the low fracture strain of these metals, fabrication of bendable and stretchable electronics from these inorganic materials is still a challenge.

However, due to the low fracture strain of these metals, fabrication of bendable and stretchable electronics from these inorganic materials is still a challenge. Conductive polymers including poly(3,4-ethylenedioxythiophene): polystyrene sulfonate (PEDOT:PSS) and functional polymers including polyaniline are used for the realisation of inkjet printing of flexible electronics [6].Conductive tracks with sinusoidal wave patterns were previously reported to offer fully reversible stretchable/compressibility without substantial strains [2]. In this study, silver and poly(3,4-ethylenedioxythiophene) polystyrene sulfonate (PEDOT:PSS) tracks were printed on an inkjet printed elastomeric substrates. The stretchability and flexibility of the1513straight line pattern and the sinewave pattern were explored and compared for both the silver and PEDOT:PSS inks.2.

The stretchability and flexibility of the1513straight line pattern and the sinewave pattern were explored and compared for both the silver and PEDOT:PSS inks.2. Materials and Methods 2.1.Materials and EquipmentObjet TangoBlack® FLX973 was used to fabricate flexible substrates in a Stratasys Connex 260 digital materials printer. The nano-silver ink (28 – 33 weight % of silver) used in this study to print conductive tracks was purchased from Advanced Nano Products (ANP). A conductive polymer, poly(3,4-ethylenedioxythiophene) polystyrene sulfonate (PEDOT:PSS) was purchased from Heraeus, Germany. Fujifilm Dimatix Materials Printer (DMP-2831) was used to print conductive tracks on the inkjet-printed specimens. Convection oven used in this study was supplied by Lenton Furnaces & Ovens. Nikon optical microscope was used to characterise the printed patterns. A TENMA (72-7725) digital multimeter was used to determine the electrical resistance of the printed conductive tracks.

Nikon optical microscope was used to characterise the printed patterns. A TENMA (72-7725) digital multimeter was used to determine the electrical resistance of the printed conductive tracks. A mechanical stylus profilometer was used to characterise the thickness of the printed conductive tracks.2.2.MethodsTensile bars were designed using Magics 14.1 computer aided design (CAD) software according to type 1 of the British Standard BS ISO 37:2011. The dimensions of the bars are given below in Table 1. For flexibility studies, cuboidal samples of 20 mm x 50 mm x 2 mm were designed in Magics. These CAD design was replicated to fabricate the required number of samples in a Stratasys Connex 260 digital materials printer. The patterns were printed using TangoBlack® with glossy finish.

These CAD design was replicated to fabricate the required number of samples in a Stratasys Connex 260 digital materials printer. The patterns were printed using TangoBlack® with glossy finish. On fabrication, the tensile bars were removed from the build platform and cleaned using deionised water.Table 1 Dimensions for the tensile bars (BS ISO 37:2011)DescriptionDimensions in mmOverall length Width of ends Length of narrow portion Width of narrow portion Transition radius outside Transition radius inside115 25 33 6.2 14 25Two strain gauge patterns (straight line and sinewave pattern) of track width 2 mm were designed using GIMP open source software. The images were designed with the resolution of 846.666 dpi. The file was exported as bitmap file (.bmp) and converted to .ptn file using the DMP800 software. A tensile bar was aligned in the build platform and the print origin was selected so that the conductive track is in the centre of the test length.2.2.1.

A tensile bar was aligned in the build platform and the print origin was selected so that the conductive track is in the centre of the test length.2.2.1. Printing of silverSince the silver ink used in this study was highly stable, all nozzles worked and hence, all 16 nozzles were used to print the conductive tracks. The print head angle was set to 6.8° to achieve a drop spacing of 30 µm. Print height was 1 mm and firing voltage was 27 V. The cartridge temperature was set to 30°C. After printing one layer of silver, the ink was dried using a heat gun (blows hot air) from approximately 15 cm above the substrate. This was performed in order to dry the excess solvents, so that the next layer can be printed. It1514should be noted that this procedure was not performed for the sintering of the silver nano- particles. After printing 10 Layers, the tensile bars containing the printed conductive tracks were transferred to an over operating at 130° C.

After printing 10 Layers, the tensile bars containing the printed conductive tracks were transferred to an over operating at 130° C. After 2 hours, the tensile bars were transferred from the oven to room temperature (~ 20° C) and allowed to cool before further characterisation.2.2.2. Printing of PEDOT:PSSThe TangoBlack® substrate for printing the conductive track was placed on top of a heated substrate. The surface temperature of the TangoBlack® substrate was maintained to ~ 50°C. The print head angle was set to 6.8° similar to silver to attain a drop spacing of 30 µm. Working nozzles were selected and print origin was selected. Print height was set to 1 mm and the firing voltage range was 19 - 23 V. No heating was applied to the cartridge and the head temperature was maintained at 20°C. After printing 10 layers, the substrate was removed and allowed to cool before further characterisation.

No heating was applied to the cartridge and the head temperature was maintained at 20°C. After printing 10 layers, the substrate was removed and allowed to cool before further characterisation. It should be noted that no further heating/blowing of hot air was performed during or after printing.2.3.Surface CharacterisationNikon optical microscope was used to characterise the morphology of the inkjet- printed TangoBlack® (tensile bar) substrates, silver and PEDOT:PSS patterns. A mechanical stylus profilometer was employed to estimate the thickness of the inkjet printed silver and PEDOT:PSS layer.2.4. Surface wettability measurementsSurface wettability of the silver and PEDOT:PSS inks on TangoBlack surface was obtained using KRUSS Drop Size Analyser (DSA) 100. The drop volume was 2 µL. The droplet formed on the substrate was allowed to settle for ~ 5 seconds before measuring the contact angle.

The drop volume was 2 µL. The droplet formed on the substrate was allowed to settle for ~ 5 seconds before measuring the contact angle. Measurements were performed on three different TangoBlack® substrates and the obtained contact angle values were averaged.2.5. Stretchability and FlexibilityAn Instron tensile machine was used to stretch the tensile bars. The tensile bars printed with conductive tracks were stretched at a rate of 1 mm per minute. Resistance was measured after every 1 mm elongation until the track stopped conducting. For flexibility measurements, cylinders offering various radius of curvature ranging from 160 mm to 7.5 mm were chosen. Samples were placed flat on the surface of the cylinder and the resistance offered by the conductive tracks was measured.2.6.Electrical ResistanceElectrical resistance was measured using a TENMA (72-7725) digital multimeter. The probes were gently placed on the contact pads at both ends of the track.

The probes were gently placed on the contact pads at both ends of the track. The ranges used were 2 Ω to 2 MΩ and the resistance values given by the multimeter was noted.2.7.Statistical AnalysisA total of three samples were used for each experiment and the results are the average of these three samples. The reported error bars and ± value corresponds to the standard error of the mean all through this study.15153. Results 3.1.Surface morphology of TangoBlack®Figure 1 shows the surface morphology of the inkjet printed TangoBlack® surface with glossy finish. A wavy pattern can be observed and is likely due to jetting and curing of the ink in swathes. The clean glossy finish of the TangoBlack® substrate was observed to be highly sticky to the substrates including glass slide and the Dimatix build platform (platen).Figure 1 Surface morphology of the inkjet printed TangoBlack® substrate obtained using an optical microscope3.2.Silver on TangoBlack® 3.2.1.

Droplets CharacterisationFigure 2 shows the ANP silver ink droplets ejected from the nozzles and the droplets formed on the inkjet printed TangoBlack® substrate. The measured diameter of the droplets using the inbuilt Dimatix software was approximately 47 ± 3.5 µm. The droplets were mostly observed to be uniform; however, in certain areas a satellite drop was observed next to the main droplet. This is possibly due to the fact that due to the wavy surface profile, the time of flight of the droplet might have been reduced. As a result, coalescence of the satellite with the main droplet might have been affected and thus leading to satellite formation next to the droplet.Figure 2 Silver droplet ejection from a Dimatix printer cartridge (left) and subsequent drop formation on the TangoBlack® substrate (right)Figure 3 shows the contact angles of the ANP silver ink droplets formed on the TangoBlack® obtained for three different samples.

This reveals that the ANP silver ink is1516hydrophilic and thus wets the surface. The wettability of silver ink on previously printed silver tracks revealed a highly wettable nature with contact angle > 20°.Figure 3 Surface wettability of the silver nanoparticle ink on the inkjet printed TangoBlack® substrates3.2.2. Characterisation of Silver TracksSurface CharacterisationOptical microscopic images of the silver tracks printed on the TangoBlack® surface (before and after sintering) has been shown in Figure 4. It can be observed from the figure that there were very few cracks for the printed silver tracks before sintering. Whereas after sintering, cracks emerge after sintering and this is possibly due to varied thermal expansion behaviour of the TangoBlack® and Silver. The presence of these cracks on the printed silver track may have a strong influence on the resistance offered by these tracks.

The presence of these cracks on the printed silver track may have a strong influence on the resistance offered by these tracks. The thickness of the printed silver tracks was estimated to be approximately 600 nm.Figure 4 Inkjet printed silver tracks on the TangoBlack® substrates (straight line and sinewave patterns) before and after oven sintering at 130°C1517StretchabilityFigure 5 shows the conductive tracks (both straight line and the sinewave pattern) printed on test specimens. These conductive tracks printed on the TangoBlack® substrates were stretched using an Instron tensile machine at the rate of 1 mm/min with 1mm steps. The resistance offered before stretching by straight line tracks were 1.1 ± 0.1 Ω and the sinewave tracks were 2.4 ± 0.2. When stretched for 1 mm, no conductivity was observed on both the straight line and sinewave track patterns; whereas after relaxing the tensile bars, the tracks were observed to conduct with slightly a higher resistance.

This shows that the cracks were formed due to stretching of the specimens.Figure 5 A photographic image showing the inkjet printed silver tracks (straight line and sinewave track patterns) on the tensile bar specimensOn examining the surface morphology of the elongated/stretched tracks through optical microscope, it was confirmed that the loss of conductivity after stretching was primarily due to the origination of cracks due to stretching (Figure 6).Figure 6 Cracks formed on the silver tracks after stretching. (a) Straight line pattern and (b) Sinewave patternFlexibilityThe substrates printed with silver were flexed to different radius of curvatures ranging from 160 – 7.5 mm. Table 2 shows the resistance obtained for the printed silver tracks on the TangoBlack® substrates.

Table 2 shows the resistance obtained for the printed silver tracks on the TangoBlack® substrates. It can be observed from the table that the sinewave pattern was capable of conducting at a lower radius of curvature than the straight line pattern.1518Table 2 Resistance measured for straight line pattern and sinewave pattern at various radius of curvaturesFigure 7 and 8 depicts the cracking behaviour of the straight line and the sinewave pattern of the printed silver tracks on the TangoBlack® substrate respectively. It can be observed from the figures that the straight line pattern was more prone to cracks than the sinewave pattern. This further confirms the ability of the sinewave pattern to withstrand for low radius of curvature compared to the straight line pattern.Figure 7 Cracks formed on a straight line silver track when flexedFigure 8 Cracks formed on a sinewave silver track when flexed15193.3.PEDOT:PSS on TangoBlack® 3.3.1.

Droplets CharacterisationDroplets formed by the PEDOT:PSS ink on the TangoBlack® substrates were almost in-visible due to its transparency. Hence it was hard to characterise the droplets for size in Dimatix. Figure 9 shows the contact angle formed by the PEDOT:PSS on TangoBlack® substrates. Wettability studies showed that PEDOT:PSS ink was wettable on the TangoBlack® substrates yielding a contact angle of 47° ± 1°. PEDOT:PSS droplets also exhibited a wettable nature (~ 35°) on PEDOT:PSS substrate.Figure 9 Wettability of the PEDOT:PSS ink on the TangoBlack® substrates3.3.2. Characterisation of PEDOT:PSS TracksSurface characterisationFigure 10 shows the surface morphology of the thin film of PEDOT:PSS printed on the TangoBlack® surface. A wavy profile can be observed and this is likely to be due to the waviness of the substrate itself. However, the printed swathes of PEDOT:PSS can also add to this waviness.

A wavy profile can be observed and this is likely to be due to the waviness of the substrate itself. However, the printed swathes of PEDOT:PSS can also add to this waviness. The layer thickness of the film was measured to be approximately 2 µm.Figure 10 Surface morphology of a thin film of PEDOT:PSS printed on a TangoBlack® substrateStretchabilitySimilar to silver on TangoBlack® substrates, printed tracks of PEDOT:PSS on TangoBlack® substrates were stretched in Instron tensile machine. Resistance measured at various extensions were graphically shown in Figure 11. As it can be observed, both PEDOT:PSS patterns withstood to about three millimetres of extension before cracking. However, the tracks conducted back once relaxed, but with a higher resistance (10 – 15 KΩ) compared to the resistance measured before elongation (1 – 2 KΩ).

However, the tracks conducted back once relaxed, but with a higher resistance (10 – 15 KΩ) compared to the resistance measured before elongation (1 – 2 KΩ). Optical microscopy images of the tracks before and after extension were presented in Figure 12.1520Figure 11 Resistance measured using an Ohmmeter for the straight line and sinewave tracks at different extensionsFew cracks on the straight track patten can be wittnessed before elongation. These cracks prossibly originated due to the sticky nature of the TangoBlack® while transfering the substrates with the printed tracks from the Dimatix printer to microscope for characterisation. However, after elongation to 4 mm, a series of cracks all through the straight line pattern can be observed. When this is compared with the sinewave pattern, cracks originated exactly perpendicular to the direction of the applied force.

When this is compared with the sinewave pattern, cracks originated exactly perpendicular to the direction of the applied force. No cracks were visible in the other areas of the printed tracks.Figure 12 Optical microscopy images showing the straight line and sinewave pattern printed using PEDOT:PSS before and after elongation. Red circles points the cracks originated in the printed sinewave track.FlexibilityThe printed PEDOT:PSS tracks were flexed to different radius of curvatures ranging from 160 mm and 7.5 mm (until the tracks stopped conducting). Table 3 shows the resistance1521obtained for the tracks flexed at various radius of curvatures. It can be noted from the table that both the straight line and sinewave patterns had same resistance within their sets when flexed until a radius of curvature of 15 mm. When flexed beyond this point, the resistance offered by the tracks increased slightly and subsequently cracked at the radius of curvature of 7.5 mm for both patterns.

When flexed beyond this point, the resistance offered by the tracks increased slightly and subsequently cracked at the radius of curvature of 7.5 mm for both patterns. Figure 13 shows the optical microscopic image of the previously flexed straight line PEDOT:PSS tracks. Cracks can be noted all through the track. Some of these cracks completely cuts the track and some are intermediate. If when flexed more, these intermediate crack may propogate further to completely cut the track.Table 3 Resistance offered by the straight line and sinewave patterns printed using PEDOT:PSS when flexed to various radius of curvaturesFigure 13 Optical microscopy image showing the cracks that originated in the printed PEDOT:PSS straight line tracks when flexedFigure 14 shows the cracks propagated for the sinewave pattern after flexing. Similar to the stretched sample in the tensile machine, cracks originate at the points perpendicular to the applied force.

Similar to the stretched sample in the tensile machine, cracks originate at the points perpendicular to the applied force. The zoomed-in image (below) shows both complete and intermediate cracks in the pattern.1522Figure 14 Optical microscopy image showing the cracks that originated in the printed PEDOT:PSS sinewave tracks when flexed. Red circles points the cracks originated in the printed sinewave track as a result of flexing.4. DiscussionSeveral studies reported inkjet printing of both conductive silver and PEDOT:PSS on various flexible substrates terepthalate (PET), polyethylene napthalene (PEN), parylene and polyimide for various sensing applications [7]–[10]. TangoBlack® is a flexible and stretchable (to a certain extent) elastomer that could be printed in Objet machine.

TangoBlack® is a flexible and stretchable (to a certain extent) elastomer that could be printed in Objet machine. Hence in this study, conductive inks were printed on a 3D inkjet printed TangoBlack® substrates with a view of printing flexible and stretchable electronics using one process ie., ink jetting.including polyethyleneTangoBlack® substrates fabricated in the Objet® machine showcased a wavy surface and obtaining a microscopically smooth surface without waviness is almost imposible in this jetting machine. However, the impact of this wavy profile on the printed surface can be reduced by choosing appropriate part orientation. In the present study, the requied substrates (both tensile bars and cuboidal samples) were printed parallel to the printing direction to minimise the wavy profile. Printing parts perpendicular to the printing direction will increase the the number of swathe per print layer, potentially leading to a highly wavy profile.

Printing parts perpendicular to the printing direction will increase the the number of swathe per print layer, potentially leading to a highly wavy profile. Printing conductive tracks on a highly wavy surface may lead to uneven printing and inturn affect the conductivity.Previous literature suggested oven sintering of silver at 130°C - 200°C for 2 hours to achieve highly conductive tracks [11]. Although 130°C was used as the sintering temperature for silver nanoparticles in this study, oven sintering of the printed silver tracks was observed to induce cracks in the printed tracks. This is most likely due to different thermal behaviour of the silver and the TangoBlack®. As it would be expected, both the straight line and sinewave patterns of printed silver tracks did not survive on elongation; however when flexed, they remained conductive for upto a certain radius of curvature.

Survival of the sinewave pattern for a lower radius of curvature compared to the straight line pattern reveals that, in addition to the material properties, track geometry also plays a part in yielding a flexible conductive tracks. Thus, optimisation of the track geometry can offer a highly conductive and flexible circuitry. Though silver tracks printed on TangoBlack® substrates may not be useful for stretchable electronic devices, it has the potential to be used in flexible electronics. However, the flexibility of these printed silver tracks are limited.1523PEDOT:PSS, the conductive polymer on the other hand was seen promising due to its transparency, thermal stability and good electrical conductivity. The PEDOT:PSS ink used in this study was a water based colloidal dispersion. Several studies, reported inkjet printing of PEDOT:PSS as an efficient and clean choice of fabricating components [12]–[14].

Several studies, reported inkjet printing of PEDOT:PSS as an efficient and clean choice of fabricating components [12]–[14]. In reality, for inkjet printing, there are issues with this colloidal dispersion of PEDOT:PSS in water. The PEDOT:PSS ink (Clevios®) supplied by Heraeus was not stable in the Dimatix print head. Although most of the nozzles (more than 12 out of 16) worked when the ink was filled, after printing few layers, the nozzles died one by one leaving with less than 4 working nozzles. If the printing was stopped for an hour or two, these nozzles were also observed to die. Ultimately, leaving the Dimatix catridge with PEDOT:PSS ink overnight led to clogging up of all nozzles. Repeated puring and wiping of the nozzles, and cleaning of nozzles with deionised water did not help to get all nozzles to work.

Repeated puring and wiping of the nozzles, and cleaning of nozzles with deionised water did not help to get all nozzles to work. This is an alarming issue and a major set-back for the inkjet printing of this water based PEDOT:PSS ink, especially in the Dimatix.Admists of these issues, the tracks printed with PEDOT:PSS showcased resistance in kilo ohms, an order of magnitude of three compared to silver ink (usually in ohms). Due to the increased track length, the sine wave pattern was more resistive (~ 2.2 KΩ) and nearly double that of the straight line pattern (~ 1 KΩ). Both the printed PEDOT:PSS patterns remained conductive upto an elongation of 3 mm and to the radius of curvature of 10 mm when flexed. PEDOT:PSS is certainly a promising material for the flexible electronics; however, the stability issues are yet to be solved. Be it a printed silver or PEDOT:PSS track, cracks were observed to originate perpendicular to the direction of the applied force and propogate further.

Be it a printed silver or PEDOT:PSS track, cracks were observed to originate perpendicular to the direction of the applied force and propogate further. Designing smart patterns such as the sinewave pattern shown in this study, origination of cracks can be controlled and limitted thereby the performance of the printed patterns can be enhanced. In contrast to straight line tracks, modification of the conductive tracks in micron scale can add more value to the printed electronics.5. ConclusionThe use of TangoBlack® as a substrate for printing silver nanoparticle ink was not promising for stretchable electronics, but has some potential for the flexible electronics. TangoBlack® can be used as substrates to print PEDOT:PSS for both stretchable and flexible electronics. However, poor stability of the PEDOT:PSS ink for inkjet printing should always be reminded. Sinewave pattern was witnessed to be a better design compared to the straight line pattern.

However, poor stability of the PEDOT:PSS ink for inkjet printing should always be reminded. Sinewave pattern was witnessed to be a better design compared to the straight line pattern. This reveals that in addition to the material property, track geometry also plays a part in determining the flexibility of the ink jet printed circuits.AcknowledgementThe authors gratefully acknowledge funding support from the Engineering and Physical Sciences Research Council (EPSRC), [UK] under the Jetting Electronic Tracks JET project, and the EPSRC Centre for Innovative Manufacturing in Additive Manufacturing (EP/I033335/2) at the University of Nottingham, UK. The authors also acknowledge the technical assistance offered by the technicians Mr. Mark East, Mr. Mark Hardy and Mr.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed Paper3D-Printable Electronics - Integration of SMD Placement and Wiring into the Slicing Process for FDM FabricationFlorens Wasserfall, Daniel Ahlers, Norman Hendrich, Jianwei Zhang Department of Informatics, Group TAMS, University of Hamburg, Germany {wasserfall, 2ahlers, hendrich, zhang}@informatik.uni-hamburg.deAbstractSeveral approaches to the integration of wires and electronic components into almost every ex- isting additive fabrication process have been successfully demonstrated by a number of research groups in the last years.

While the pure mechanical process of generating conductive wires inside of a printed object has proved to be feasible, the design, integration, routing and generation of toolpaths is still a laborious manual task. In this paper, we present a novel approach to place and wire SMDs in a three-dimensional ob- ject, based on schematics generated by conventional PCB design tools such as CadSoft EAGLE. Routing wires in an object for FDM manufacturing requires certain knowledge about the printer’s properties to meet the extruder characteristics, avoid non-ﬁllable regions and electric shorts. Cor- respondingly for the slicing of conductive wires, the software must respect appropriate channel widths, avoid interrupted traces and ensure proper endpoints serving as contact pads for the SMDs. To fulﬁll those requirements, we implemented the design and routing software as a native extension of an existing slicing software.

To fulﬁll those requirements, we implemented the design and routing software as a native extension of an existing slicing software. The user works in a three-dimensional representation of the ﬁnal extruder toolpath, augmented by the routing information. The actual computing step is executed at the layer level by manipulating the polygons which represent the two-dimensional object topology and toolpath for each single layer, allowing the routing algorithm to avoid the generation of non- printable traces. We successfully designed and printed some test objects including a force-sensor prototype, demonstrating a signiﬁcant improvement in the usability and efﬁciency over manual solutions.1. IntroductionThe integration of electric connections and electronic components into additively manufactured parts is considered a key feature for the application in various ﬁelds such as wearable devices, (soft)-robotics or prototyping applications.

The desire for printing conductive materials has been addressed with an increasing number of different approaches over the past decade. Direct writing of conductive, polymer based silver inks has been successfully demonstrated for SLA [1, 2], and FDM [3, 4, 5] printed objects and free-form surfaces [6]. The ink is dispensed through a nozzle directly to the object’s surface from a pneumatic pump or a syringe. Embedding of wires or ﬁbres is a second common technique where a copper wire is either injected into the printed plastic surface by a heated nozzle [7] by ultrasonic welding [8, 9] or by encapsulation into the ﬂuid plastic immediately after extrusion [10]. Other approaches explore the use of inkjettable inks [11, 12, 13] or carbon ﬁlled ﬁlament [14].

Other approaches explore the use of inkjettable inks [11, 12, 13] or carbon ﬁlled ﬁlament [14]. A recent comprehensive review is given by [15].1826Most of the work mentioned above is still limited to simple proof of concept applications where the circuit design and toolpath generation can be done manually or with simple scripts. The in- tegration of circuits with higher complexity clearly requires appropriate design tools. A simple solution for this problem is the manual design of channels and cavities with common CAD-tools e.g. Solidworks or OpenSCAD. The object is then printed on a normal printer and the conductive traces are applied in a post processing step [3, 1] or exported as a multi-material model which can be printed directly by a dual-extruder printer [5]. Bayless et al. used a MatLab script to gener- ate G-code from parametric curves and rectilinear patters [7]. It is not stated how the curves and patterns are generated. Macdonald et al.

Bayless et al. used a MatLab script to gener- ate G-code from parametric curves and rectilinear patters [7]. It is not stated how the curves and patterns are generated. Macdonald et al. introduced the concept of deforming ﬂat circuit designs created with common PCB design tools [16]. They demonstrate applications where a ﬂat circuit is wrapped around a cylindrical object and an “unfolded” cube which represents the six sides of a gaming dice. In 2015, Swensen et al. presented the concept of computing a system of tubes with varying diameters from a 2-dimensional board layout created in EAGLE [17]. Through-hole com- ponents are inserted through connected holes and the tubes are ﬂooded with injected low-melting metal alloy. The ﬂow is controlled by varying friction determined by the tube diameter and length to reach every pin simultaneously. The commercial solution Project Wire [18] was announced by Autodesk and Voxel8 and is currently available in a public alpha version.

The commercial solution Project Wire [18] was announced by Autodesk and Voxel8 and is currently available in a public alpha version. A set of predeﬁned electronic components can be integrated into STL-ﬁles and connected via traces in a web-based drag-and-drop application. The result is exported as two separate meshes, one representing the electric wires, the other representing the difference of the original object and the wire traces.This paper introduces the concept of a design software, aiming to close the gap between available printing processes and common CAD tools for object and circuit design. It is implemented as an extension of the open source slic- ing software Slic3r [19]. The general idea is to load an existing object mesh and a schematic created by a common PCB design tool and place components at different layers of the sliced object, similar to the layout step for a PCB board as pictured in Fig. 1.

1. The user can vertically “browse” through the ob- ject and place components at arbitrary positions inside the object or at its sur- face. Unrouted connections are visual- ized as rubberbands. High level rout- ing is achieved by setting waypoints, which serve as interpolation points for the low level routing to generate extru- sion paths, approximating the desired route as close as possible while opti- mizing physical printability.Figure 1 – Electronic components embedded into a cus- tom 3D-object for FDM manufacturing, showing the wiring layer of a humanoid robot (shoulder) ball-joint with integrated position sensing. Visible are the 3-axis hall- effect sensor at the center of the ball joint (blue, lower left), the microcontroller (blue, near the center), and some pas- sive SMD components. Our slicing tool shows the user- deﬁned wiring waypoints (red) and automatically calcu- lates the wiring extrusion (light blue), channel boundaries (yellow) and inﬁll (purple).

Our slicing tool shows the user- deﬁned wiring waypoints (red) and automatically calcu- lates the wiring extrusion (light blue), channel boundaries (yellow) and inﬁll (purple). Also shown is the printbed ad- hesion layer (green) required for FDM manufacturing with ABS material.18272. Design processManufacturing an integrated 3-dimensional circuit generally requires a model of the object, a cir- cuit design and a transformation into an executable toolpath (slicing). The integration of model and circuit could potentially be implemented into one of the corresponding tools: CAD software, PCB design tool or slicer, or as an additional standalone application.

The integration of model and circuit could potentially be implemented into one of the corresponding tools: CAD software, PCB design tool or slicer, or as an additional standalone application. This section motivates and explains our design decisions for the integration of all three aspects.3D-Electronics design approachesCreating 3D-electronic designs manually with a CAD software is only feasible for simple test ob- jects since it is laborious, very time consuming and has no reference to any electronic component data.A more promising approach is the creation of PCB layouts with a PCB designer and either deform- ing or combining them with a CAD software or to use multilayer PCB circuits with a subsequent registration step that maps each layer of the PCB to a corresponding slice. Both methods can bene- ﬁt from the well-engineered PCB design software and use features like the existing component databases, electronic rule checking and simulations.

Both methods can bene- ﬁt from the well-engineered PCB design software and use features like the existing component databases, electronic rule checking and simulations. However, this approach does not produce real 3D electronics because the ﬁnal circuit is still limited to a combination or deformation of 2D PCB designs.Standalone 3D electronic design solutions can simply design real 3D electronics without limita- tions for component positions and wire paths. They are very ﬂexible due to their independence from surrounding tools like the slicer or the CAD tool. A problem that all these approaches have in common is the missing reference to the actual structure of the print. Without this reference, the designer cannot react to structural differences that 3D printed objects have and the slicer is not able to handle wires and components differently.

Without this reference, the designer cannot react to structural differences that 3D printed objects have and the slicer is not able to handle wires and components differently. If slicing information is available, it can be used to ensure a predeﬁned wire thickness, to prevent shorts, to create proper SMD pads, and to ensure the connection on intersections of wires without overﬁlling them. Project wire is currently the only implementation of this approach. However, it has no schematic design feature; the designer has to create the circuit connections while he places the components inside the object. This can lead to missing connections or wrongly connected components.Implementation considerationsSince all the existing 3D electronic design approaches have some advantages, the new process tries to combine the advantages of the existing approaches. A PCB design software is used for the schematic design, enabling us to use the existing functions and the component database.

A PCB design software is used for the schematic design, enabling us to use the existing functions and the component database. However, the PCB design software is incapable of designing actual 3D electronic circuits. Therefore, plac- ing of components and wire routing is implemented as an extension to an existing slicing software. During the slicing step, all relevant information is available: the CAD- and PCB-Data and the ac- tual structure of the resulting print.1828The proposed process (Fig. 2) starts with the speciﬁcation of the desired device. Then the CAD design and the schematic design of the circuit can be done simultaneously. The 3D-model and the schematic are imported into the slicing tool. The slicer repeatedly runs through the slicing cycle every time a layer changes. The changes in the layers are triggered by adding components or wires. The cycle starts with the slicing of the whole model and generating the structures for each slice, beginning with the bottom slice.

The cycle starts with the slicing of the whole model and generating the structures for each slice, beginning with the bottom slice. The slicer ﬁrst generates the perimeter of the object. Then the cavities for the components are generated if they in- tersect the slice. Wires and wire chan- nels are generated if they are in the slice. Lastly the remaining space is ﬁlled with inﬁll structure. When the design is ﬁnished the G-code can be exported. The modiﬁed slicer is designed to handle different schematics ﬁle formats. Due to its simple XML ﬁle format, EAGLE was used as a PCB design tool. To use another PCB design software, a simple converter has to be written that converts the schematic ﬁle format to the internally used format. Additional information regarding the formats can be found in previous work [20]. To save the design, only the additional information of the placed components and the wires is stored in a new ﬁle.

Additional information regarding the formats can be found in previous work [20]. To save the design, only the additional information of the placed components and the wires is stored in a new ﬁle. The information which is extracted form the schematic ﬁle is just referenced in the saved ﬁle. This ﬁle structure allows editing the schematic with the PCB design tool in the whole process.Specification3D objectElectronics schematicSlicingPerimeter generationAdd components/ route wires (designer)Infill generationWire generationCavity generationG-code exportFigure 2 – Overview of the iterative design cycle for 3D-printable electronics3. Embedding of SMD-componentsA 3D-model of each component is required to visualize and place the components in the object. Since EAGLE does not provide those 3D models, they have to be generated during the import step. To generate the 3D models, the outline of each component and the footprint is expanded to a 3D model.

Since EAGLE does not provide those 3D models, they have to be generated during the import step. To generate the 3D models, the outline of each component and the footprint is expanded to a 3D model. The height of the footprint relies on the layer thickness from the sliced model to ensure a proper connection to the ink trace. The height of the component is set to a predeﬁned value that can be set individually for each package. A database of standard package dimensions would simplify this step for the user and could also help render complex components. To place a component, the designer has to select a layer of the object and place the part on this layer. The component can then be rotated around all three axes.

To place a component, the designer has to select a layer of the object and place the part on this layer. The component can then be rotated around all three axes. A cavity for each component is generated by ﬁrst computing a polygon of the component’s shape, offsetting this outline by a user deﬁned amount to provide some tolerance and ﬁnally subtracting the polygon from the layer, which is also represented by a number of nested polygons [21, 22]. This step is repeated for every layer intersecting the current component (Fig. 3.1).1829Figure 3 – Subtraction of SMD components from the printed object to create cavities. From left to right: 1: outline subtraction, 2: part rotated but subtracted outline generated from ﬂat footprint, 3: naive polygon subtraction with rotation, 4: accumulating polygon subtraction with rotation.The naive outline subtraction cannot handle rotation around the x or y axis (Fig. 3.2).

3.2). To handle rotated components, the polygon of the rotated component has to be computed and subtracted from the object (Fig. 3.3). For components which are not located at the surface of the object, the cavity must be extended along the z axis to allow placing of the component without extruder collisions. This can be achieved by subtracting the union of all polygons from intersections with layers below the current layer (Fig. 3.4). With this extension it is possible to insert the component into the object directly before the cavity is covered.4. Wire generationAll wires connecting the individual components are organized as nets by the PCB design tool. Each net contains all pins with equivalent electrical potential and represents the physical connections between this speciﬁc subset of pins. In our tool, the “rat’s nest” of placed but currently unwired components is visualized by a set of rubberbands.

In our tool, the “rat’s nest” of placed but currently unwired components is visualized by a set of rubberbands. Each net is represented by a set of points and a set of connections between these points, forming an acyclic undirected graph. Rubberbands, points and parts are represented in the GUI as active elements to provide user interactions. The resulting wires are visualized as extrusion objects embedded into the toolpath representation but highlighted by a different color (compare Fig. 4).Routing conceptsThe routing of wires through the 3-dimensional object is divided into two steps: the high-level generation of “waypoints” to deﬁne the desired routing paths and the low-level routing of a single line between each tuple of waypoints.

This concept allows for a certain level of decoupling be- tween the classic routing task of ﬁnding a collision free partitioning of the available space given a number of constraints, and the process related task of ﬁnding a set of suitable extrusion paths given a speciﬁc printer’s parameters. The waypoints are currently manually deﬁned by the user by manipulating the rubberbands in the 3D-visualization. This can easily be extended to import a set of waypoints from an external 3D-electronics design tool or a routing algorithm.Low level routingThe low level routing step computes a set of extrusion skeletons for each layer from the 3-dimen- sional set of waypoints. This process is similar to the slicing step which divides an object into a stack of 2.5-dimensional layers by computing the intersection of a set of planes with the tesselated mesh representation.

This process is similar to the slicing step which divides an object into a stack of 2.5-dimensional layers by computing the intersection of a set of planes with the tesselated mesh representation. For a given layer L and each wire segment (waypoint tuple) [a, b], the inter-1830Figure 4 – Prototype of a 2DOF force sensor. top-left: EAGLE schematic of the circuit consisting of a microcontroller and two optical proximity sensors. top-right: Upper slicing layer, the top of the two optical sensors and the microcontroller and their cavities are visible. bottom-left: Middle slicing layer containing most of the components and wires. Waypoints (red) and conductive wires (purple) are visible. bottom- right: Lower slicing layer, unwired (black) and wired (gray) rubberbands are also displayed on this layer.

Waypoints (red) and conductive wires (purple) are visible. bottom- right: Lower slicing layer, unwired (black) and wired (gray) rubberbands are also displayed on this layer. Two wires have been moved to this layer to avoid intersections.section with the lower and upper boundary Lb, Lt is computed bya(cid:2) x=⎧ ⎨⎩ax + (bx − ax) · Lb−az bz−az ax + (bx − ax) · Lt −az bz−azif az < Lb if az > Ltand correspondingly for a(cid:2) y and point b, resulting in a 2-dimensional set of lines. This step is trivial for horizontal wire segments, where the intersection equals the planar projection of the original wire. For inclined wires, the start point of the resulting line is then extended by a user deﬁned offset to achieve a stable inter-layer connection to the adjacent wire section if a is outside the boundaries of L.Extrusion pathsThe result of the low level routing step is stored as an unordered vector of lines (as depicted in Fig. 5).

5). Each extrusion path that can be printed as a continuous line is assembled by:1831Figure 5 – Generation of extrusion paths for a single layer. From left to right: 1: Wire segments from slicing. 2: Assembling a connected graph from unordered segments. 3: Re-fragmentation at junctions, thin black arrows depict directions of overlapping extrusions. 4: Printed layer. Overlapping regions to ensure extrusion towards the SMD-pads are marked in red.1. Finding an endpoint. Iterate the vector until the ﬁrst line with at least one endpoint which doesn’t coincides with any other endpoint. Remove this line from the vector.2. Traversing adjacent lines until ﬁnding a point which coincides with 0 (endpoint) or more than 1 (intersection) other points and removing the lines.3. Repeat steps 1-2 until the vector is empty.The resulting polylines are used as a basis for all subsequent steps.

Repeat steps 1-2 until the vector is empty.The resulting polylines are used as a basis for all subsequent steps. Channels for the wires are generated by offsetting the polylines by e + Δ and subtracting the resulting polygon from the current 2 layer. e is the extrusion width of the conductive material extruder, Δ is an offset value provided by the user.Figure 6 – Different layers of a toolpath with integrated electronic components and wires, generated by the augmented slicing tool. Layers in ascending order from left to right: 1: Internal layer with sparse inﬁll (red lines, density 50%). 2: Layer directly below wires and components. The internal perimeters (yellow) serve as a smooth “substrate” for the wires. The inﬁll density is increased at regions where components will be placed in the next layers, to provide a solid surface, covering the sparse inﬁll. 3: Channels and cavities are generated for the extrusion of wires and components.

3: Channels and cavities are generated for the extrusion of wires and components. 4: Top layer with placed components.1832The polylines can be directly used as trajectory for the direct write extrusion. However, since the starting point of a an extrusion path tends to be not reliable after a retraction move, each trajectory is split into two segments and the extrusion is executed from the center of the line to the SMD-pad. The segments are extended by a user deﬁned offset extrusion overlap to ensure a proper electrical connection. Possible over- or underextrusion only occurs at non-critical spots in the object. Priming the extruder is most important after a toolchange. Hence the segments are ordered by length and the amount of overlapping is increased only for the ﬁrst segment of each layer to minimize double extrusions.

Hence the segments are ordered by length and the amount of overlapping is increased only for the ﬁrst segment of each layer to minimize double extrusions. As a ﬁnal step, “substrate” is generated for the next layer by offsetting the polylines by a small value ε and subtracting the resulting polygon from the contour of L−1. This causes the extrusion of a perimeter, where the “gap” is precisely aligned with the wire to avoid disruptions of the wire extrusion due to a crossing, rough surface as described in [8].Performance considerationsSince the user is working in a representation of the ﬁnal toolpath, every interaction e.g. moving a part, routing a wire requires re-slicing and generation of the entire toolpath. For complex objects, this can easily take seconds up to minutes, resulting in poor user experience and high latencies. To address this issue, the layer representation is extended by a boolean ﬂag, indicating whether this layer is affected by a certain user operation.

To address this issue, the layer representation is extended by a boolean ﬂag, indicating whether this layer is affected by a certain user operation. The toolpath generation steps are then only executed for those layers marked “dirty”.5. Evaluation / experimental resultsAll experiments were conducted by a modiﬁed K¨uhling&K¨uhling RepRap Industrial FDM-printer [23]. The general design of the modiﬁcations is an evolution step from the CNC-mill based pro- totyping machine which was introduced in our previous work [5]. The printer is equipped with an additional syringe extruder for conductive ink, a pivot-mounted vacuum gripper and two camerasFigure 7 – Modiﬁed RepRap Industrial printer, augmented by a conductive ink extruder, vacuum gripper for component placing and cameras to align components.1833Figure 8 – Fabrication of the 2-DOF force-torque sensor prototype shown in Fig. 4.

4. The device uses infrared proximity sensors to measure the deﬂection of deformable beams with known stiffness and computes the force vectors on an embedded 8-bit microcontroller.for robotic placement of SMD-components. The modiﬁcations were facilitated by the fact that both hardware design and software of the printer are open source and all parts are speciﬁcally de- signed to be printable on the machine itself.To test the algorithms with a sufﬁciently complex design, we fabricated the prototype of a 2-DOF force-torque sensor (see Fig. 4 and 8). The sensor consists of a deformable beam structure with known mechanical properties. The thickness of the beam structure is computed according to the desired measuring range.

4 and 8). The sensor consists of a deformable beam structure with known mechanical properties. The thickness of the beam structure is computed according to the desired measuring range. The beam-deﬂection induced by an applied force is measured by infrared proximity sensors which are recessed into the surface and the deﬂection distance is then converted into the applied force by the embedded microcontroller.The quality and reliability of the printed wires is generally highly dependent on the printers cal- ibration. Even very small z-offsets of the ink extruder cause interrupted extrusions and broken wires due to the polymer’s high surface tension. Fig. 9 shows an effect that occurs if the plastic extruder “jumps” crossing the channel. A small amount of plastic is drawn as a thin string from the substrate extruder’s hot-end, later causing interruptions at the ink extrusion. Those strings are virtually invisible for the eye.

A small amount of plastic is drawn as a thin string from the substrate extruder’s hot-end, later causing interruptions at the ink extrusion. Those strings are virtually invisible for the eye. This effect can be mitigated by using plastic with a higher glass transition temperature and lifting the z-axis by several millimeters during the jump.Figure 9 – Interrupted conductive ink extrusion due to thin plastic strings caused by insufﬁcient retraction of the PLA base material (left) and successful extrusion on ABS material (right).18346. Conclusion and future workThis paper presents the concept and ﬁrst implementation of a software for the integration of electric circuits into 3D-printed objects. So far, we have only tested simple circuits with a few components, but the increase in usability over existing approaches is clear, providing a new tool for researchers and developers.

So far, we have only tested simple circuits with a few components, but the increase in usability over existing approaches is clear, providing a new tool for researchers and developers. The design is modular and relies on open source projects and widely used com- mon standards and formats.The pending next development steps will focus on improvements of the user interface and routing algorithms. Our current work is dedicated to the implementation of low level routing approaches for better utilization of the available space inside the objects and alignment to existing structures to improve the quality. Important concepts include “Grid alignment” as a simple limitation of part and waypoint positions to a predeﬁned grid as commonly used in PCB layout tools. The grid resolution is determined by the diameter and spacing of extrusions and the common pitch of the SMD components.

The grid resolution is determined by the diameter and spacing of extrusions and the common pitch of the SMD components. An additional approach is to follow the contour of existing perimeters and prior deﬁned wire routes by ﬁnding an entry and exit point and offsetting the shortest polygon segment between the two points.More information, including the source code is available on the projects website: https://tams.informatik.uni-hamburg.de/research/3d-printing/ conductive_printing.References[1] Erick De Nava, Misael Navarrete, Amit Lopes, Mohammed Alawneh, Marlene Contreras, Dan Muse, Silvia Castillo, Eric Macdonald, and Ryan Wicker. Three-Dimensional Off-Axis Component Placement and Routing for Electronics Integration using Solid Freeform Fabri- cation. Proceedings of the 19th International Solid Freeform Fabrication Symposium, pages 362–369, 2008.[2] Amit Joe Lopes, Eric MacDonald, and Ryan B. Wicker.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed Paper3D-PRINTING GRAPHENE OXIDIZE BASED ON DIRECTIONAL FREEZING Feng Zhang*,Qiangqiang Zhang∆, Weston Grove†, Dong Lin† and Chi Zhou*Department of Industrial and System Engineering, University at Buffalo, Buffalo, NY †Department of Industrial and Manufacturing Systems Engineering, Kansas State University, Manhattan, KS ∆Department of Mechanics and Engineering Sciences, College of Civil Engineering and Mechanics, Lanzhou University, Lanzhou, Gansu 730000, P. R. China.AbstractThis paper aims to provide a new process that is based on micro-dispensing and directional freezing to fabricate macro and micro controllable 3D graphene aerogel.

R. China.AbstractThis paper aims to provide a new process that is based on micro-dispensing and directional freezing to fabricate macro and micro controllable 3D graphene aerogel. In the first section, a design model of the proposed system to print 3D graphene oxide is presented, and the configurations are discussed in detail. The presented new method is contrasted to other few graphene 3D printing process. A process planning is provided includes the complete fabrication process and printing process. The physics mechanism behind the process is illustrated. A list of 2.5D and 3D printed samples are shown. Graphene Oxide solution is an easy to print material for micro-dispensing device, we successfully printed GO solutions in a stable and reliable way. Our freezing based 3D printing process matches well with freeze drying technology, which together composes the key step for fabricating truly 3D graphene aerogel.1.

However, tailored macro-architecture of graphene aerogel for specific application via controllable and scalable assembly method remains a significant challenge. In order to fully exploit the exotic physicochemical properties of 2D graphene material and explore its broader range of macroscopic applications, there is an urgent need to develop a new 3D printing technique that can print pure and continuous graphene aerogel with truly 3D architecture.3D printing is a disruptive manufacturing technology that can selectively deposit material in 3D space to form complex object. This unique property makes 3D printing the best candidate for fabricating engineered graphene aerogel with complex structures. However, due to the universal challenge of material printability, it was until recently that some extrusion based 3D printing techniques have been applied to fabricate graphene 3D periodic architectures (Zhu et al., 2015, García Tuñon et al., 2015).

In both work, pure GO (graphene oxide) solution is modified in respect to the rheology behavior to match their extrusion based process. In Zhu’s work, Fumed Silica was added in the GO suspension (40 mg ml-1) to achieve highly viscous and shear thinning non-Newtonian fluids. Branched copolymer surfactant (BCS) and Glucono-δ-lactone (GδL) are added into the GO solution in Esther’s paper. However, these extrusion based 3D printing approaches suffer from several disadvantages. First, these complex GO suspension manipulation process accompanies more complicated post-process thereafter, such as etching of the silica filler in hydrofluoric acid solution, which will cause some residual of the fillers in high risk. Second,‐1611the extrusion is a continuous printing process, only simple rod-like structures are possible to print; moreover, commonly used nozzles are several hundred micron in diameter, and thus it cannot achieve the desired dimensional accuracy.

Third, a common issue for layer manufacturing is bonding between two layers, this issue is especially prominent in extrusion based process such as Fused Deposition Modeling (FDM).In this paper, a novel process for printing macro and micro controllable 3D graphene aerogel based on directional freezing technology is presented. Similar technology has been utilized to develop Rapid Freeze Prototype (RFP) 3D printing process, which is a novel Additive Manufacturing (AM) technique that can build three-dimensional ice objects directly from computer-aided design (CAD) models by depositing and rapidly freezing specific liquid droplets layer by layer. Some related research has been done based on this process (Barnett et al., 2009, Ossino et al., 2009, Bryant et al., 2003, Zhao et al., 2010, Leu et al., 2000, Liu and Leu, 2007, Sui and Leu, 2003, Bryant and Leu, 2009).

However, only water was used as printing material, which can mainly be used for visualization and molding purpose with very limited applications. In this paper, we utilize this technique to rapidly freeze the water based graphene oxide (GO) droplets into ice crystal on a cold sink substrate as low as -20 ºC. The dilute pure aqueous GO suspension not only removes the addition of silica filler which otherwise will affect the aerogel structure integrate, it also allows to fabricate lower density and larger surface area aerogel due to the low GO concentration. Unlike the continuous extrusion based 3D printing process, which suffers from bad bonding due to weak intermolecular diffusion as well as the undesirable voids between the adjacent filaments. In our developed process, when the liquid solution is deposited on top or side of the previously frozen material, it partially melts the already frozen material.

In our developed process, when the liquid solution is deposited on top or side of the previously frozen material, it partially melts the already frozen material. These two materials will be mixed together and refreeze under low temperature environment. As the re-melted aqueous material possesses low viscosity, the voids are instantly filled by the liquid material under surface tension and gravity. Since the deposited materials bind with the previous layer firmly through hydrogen bond, high structures integrity of the final assembled graphene aerogel can be achieved. Moreover, our process is based on drop-on-demand micro-dispensing technology, a nozzle with 80 micron is used to print the GO solution, which means ultra-tiny feature can be formed.

Moreover, our process is based on drop-on-demand micro-dispensing technology, a nozzle with 80 micron is used to print the GO solution, which means ultra-tiny feature can be formed. Our preliminary findings show that the proposed technique can not only control the macro-scale architecture but also manipulate micro-scale morphology with high accuracy possibility (Zhang et al., 2016).The remainder of the paper is organized as follows. In section 2, we overviewed the proposed system including the hardware configurations, software used and control system. In section 3, we proposed the process planning schematic for this presented process, and the key parameters are discussed. In section 4, we outlined the main steps to prescribe the GO solution for our printing process. In section 5, the mechanism of directional freezing is discussed. In section 6, we present the experiment results and show some finished 2.5D and 3D graphene aerogel structures.

In section 5, the mechanism of directional freezing is discussed. In section 6, we present the experiment results and show some finished 2.5D and 3D graphene aerogel structures. Finally we conclude and discuss future work in section 7.2 Experimental setup 2.1 Hardware Configuration:A Fused Deposition Modeling (FDM) based 3D printer is retrofitted to work as the host device of our project. The XY positioning precision of the printer is12 microns, and Z positioning precision is 5 microns. Figure 1 (a) shows the main components in the configuration. A peltier is1612used as the cold plate, and mounted together with a heat sink to move up one layer each time during printing. A solenoid equipped syringe is mounted on a medium that can move along XY axis to travel along designed path in the printing process. The syringe is pressurized using a Nordson Performus™ V dispenser with certain pressure.

The syringe is pressurized using a Nordson Performus™ V dispenser with certain pressure. The two solenoid contact pins are connected to the solenoid drive which is then connected with an Atmega2560 board. The physical setup of the whole system is shown in Figure 1 (b).Figure 1 Designed model and the physical modelThe Atmega2560 board is used to control the micro-dispensing solenoid from THE LEE CO. A code is developed and burn into the Arduino board to control the ON/OFF status of solenoid. Four parameters are required to input at the Arduino serial monitor window to instruct the board, including pin number to teach which pin is active, vibrating status to control whether the pin vibrates or not, frequency of the vibrating, and duty cycle of each vibrating.2.2 Software and control SystemThe motion of printer is controlled by the generated G-code, and all the travel uses linear interpolation, no circle interpolation is needed.

To coordinate the motion of printer and dispensing of solenoid value, the connection between the extra Atmega2560 Arduino board and the one in the printer needs to be built. For multi-material printing, the g-code is modified so that when it shifts between different materials, one solenoid is disabled and the other is activated. Third, for two successive layers, a waiting time is added in the code so that the layer to drop on is fully frozen.3 Process planning3.1 General fabricating processThe GO ink is controllably ejected onto a cold plate with a certain amount of flow rate and moves following a designed routing. GO ink is instantly frozen to ice after deposited onto the cold plate which is around temperature of -20oC.

GO ink is instantly frozen to ice after deposited onto the cold plate which is around temperature of -20oC. The as-formed 3D periodic graphene oxide lattice (PGOL) is then submerged into liquid nitrogen to be firstly frozen in critical cooling conditions about -190 oC after the process of 3D printing is finished, and then placed in an ultralow temperature chamber (-80 oC) for 24h to implement further ice crystallization. 3D PGOL aerogel is obtained by followed freeze drying process for 48 h. Finally, the PGOL is thermally annealed at 1000 oC for 1 h in a tube furnace under argon atmosphere to obtain periodic graphene lattice (PGL)1613with highly conductive and lightweight properties. This process is shown in Figure 2. The aqueous GO suspension and water are loaded in two reservoirs (a). Computer program converts the input 3D model into machine command and control the nozzle jetting and motion, and the materials are selectively dropped onto a cryogenic substrate (b, c).

Computer program converts the input 3D model into machine command and control the nozzle jetting and motion, and the materials are selectively dropped onto a cryogenic substrate (b, c). The 3D printed ice structure is immersed into liquid nitrogen for further self-assembly of GO sheets (e). After the ultra-low temperature treatment, the green part is frozen under -80 oC for 24 hours (f) and the water is removed by freezing and vacuum drying for another 24 hours (g). The freezing dried part is then reduced by thermal reduction (h) and the final graphene aerogel with ultra-low density is produced (i).Figure 2 Graphene Aerogel fabricating process3.2 Printing processAs a sub-section of the whole fabricating process, the graphene oxide printing process is fundamentally important.

A CAD model of the print is designed; the CAD model is converted to .STL format; the STL file is then sliced using a prescribed software with delicately tuned parameters; the generated G-code is modified specifically so that it can be recognized by our retrofitted printer; finally the G-code is uploaded to the host software for printing. This printing process is shown in Figure 3. Unlike other 3D printing process, in which support needs to be specifically designed to both minimize the material consumption and mechanically hold the part, the support material in an aqueous dispensing process is simply the negative shape of the part with regarding to its minimum bounding box. However, the ideal support material should satisfy some requirements. First, preferably similar rheology behavior is expected so that they can form unity contact angle upon dropped which is necessary to get a smooth surface.

First, preferably similar rheology behavior is expected so that they can form unity contact angle upon dropped which is necessary to get a smooth surface. Second, the support material needs to share similar freezing temperature with the matrix material so that they can freeze similarly in one set up, more ideally they have similar latent heat. Finally, the support material needs to be easily removed in post process.

Finally, the support material needs to be easily removed in post process. The water used in our process naturally satisfies all1614these requirement since the majority of our GO solution is also water, and all the water will be sublimated from solid phase to gas phase directly in the freeze drying process.CAD 3D ModelYNeed support or notNegative shape designNSTL FileSeperately export STL FilesG-code FileCombine STL FileModified G-code FileG-code FileUpload G-code File to host SoftwareModified G-code FileSingle head printingUpload G-code File to host SoftwareRemove model and postprocessingDual head printingFigure 3 Graphene oxide printing process3.3 Path planningFor a liquid deposition, a 100% infill should be the proper density otherwise liquid will flow anywhere that is not filled in previous layer. Figure 4 (b) shows a typical rectilinear infill pattern.

Figure 4 (b) shows a typical rectilinear infill pattern. Figure 4 (c) shows the path planning of a layer of model in (a) with graphene oxide only, and Figure 4 (d) shows the path planning of a multi-material layer of the model, the red-framed portion is water, and the rest is printed with graphene oxide.Figure 4 Path planning (a) A model of lattice with support (b) Rectilinear infill pattern (c) A layer with only graphene oxide (d) A layer with both graphene oxide and water16154 Material preparationGO ink Preparation: Single layer GO precursor was purchased from Cheap Tubes Inc. (USA) with average size of 300-800nm lateral dimensions and 0.7-1.2 nm layer thickness, which was synthesized by a modified Hummers’ method featured 40-60% functional group content (C: 35-42%, O: 45-55%, H: 3-5%) providing high solubility in DI Water with high purity of 99 wt.%. GO powder was ultrasonically dispersed in DI water to prepare uniform suspension with concentration of 10 mg/ml.

GO powder was ultrasonically dispersed in DI water to prepare uniform suspension with concentration of 10 mg/ml. An agent ethylenediamine (EDA) was mixture with GO suspension by volume ratio of 2‰ to prepare the 3D printing used GO ink. The mixture of GO and EDA was ultrasonically dispersed for 0.5 h to make it uniform and then added NaOH solution (Concentration= 5%) by volume ratio of 5‰ for adjusting the viscosity of ink to be suitable for ink-ejecting process. The stacking GO sheets are bonded by π-π attraction which is classified as non-covalent bonding. The weak non-covalent bonding between the GO sheets cannot fully unlock the exotic properties of the 2D graphene sheet. Our primarily results show that the chemical covalent bonding between the GO sheet surface and the reduction agent can give rise to improved mechanical, thermal and electric properties.

Our primarily results show that the chemical covalent bonding between the GO sheet surface and the reduction agent can give rise to improved mechanical, thermal and electric properties. And EDA serves as the reducing and functionalizing reagent in our process, it play an important role in the assembly of graphene oxide sheet. Figure 5 illustrates how the EDA (in the middle) bonds up and down two different GO layers.Figure 5 Schematic illustration of mechanism of EDA bonding with GO layers5 Freeze castingFreeze casting (also known as ice templating) is a technique that takes advantage of the anisotropic solidification behavior of a solvent (Generally water) with well dispersed solute (Extensively ceramic slurry) to form a directionally porous structure. The first observation of cellular structures resulting from the freezing of water can date back to more than a century ago (Lottermoser, 1908).

The first observation of cellular structures resulting from the freezing of water can date back to more than a century ago (Lottermoser, 1908). However, it wasn’t until 2001, when Fukasawa created directionally porous alumina castings (Fukasawa et al., 2001), that the idea for using freeze-casting as a means of creating novel porous structures really set sail. Since then, research has grown considerably in this area. Freeze casting has been used to fabricate metal/ceramic and model polymer/ceramic (poly(methyl methacrylate); PMMA)/alumina and, more recently, PMMA/SiC hybrid materials with fine lamellar or brick-and-mortar architectures (Deville et al., 2006, Liu et al., 2014, Munch et al., 2008, Roy et al., 2010, Wegst et al., 2014).In the situation of graphene oxide, a well-dispersed partially reduced GO dispersion is frozen, pr-GO sheets are concentrated at the boundary of ice crystals and then aligned along the growth direction of ice due to the squeezing effect.

As a result, a continuous honeycomb-like network is formed (Qiu et al., 2012, Hunger et al., 2013). The mechanism of the freeze casting to form ice crystals repelling the GO particles is illustrated in Figure 6 (a). After the well dispersed GO sheets is frozen, they are expelled by the ice crystals and aligned along the crystal growth1616direction.

After the well dispersed GO sheets is frozen, they are expelled by the ice crystals and aligned along the crystal growth1616direction. Due to π-π attraction between the GO sheets and the functionality of EDA, the assembled network was found to be very stable and maintain its structure integrity upon sublimation of ice which finally results in the formation of highly organized hierarchical porous structure (Figure 6 (c)-(d)).Figure 6 the formed micro structure of 3D printed Graphene macro structure (a) Growing ice crystals expel 2D graphene oxide particles into crevice (b) Macro image (c)-(d) The large magnification SEM image of surface morphologies with hierarchical porous nanostructures (pore size ~ 250 μm, facial-link length ~ 170 μm)6 Results and discussionsThis paper developed a novel 3D graphene aerogel fabrication process. The process is composed of two major parts.

The process is composed of two major parts. One part is 3D printing graphene oxide, we have experimented with our machine for a bunch of different structures including 2.5D network, grids, 3D lattice, turbine, gears etc. The other part is forming final graphene aerogel, for each printed sample, we carefully processed them to form final aerogel.6.1 Printed graphene ice structuresSeveral structures have been successfully built including 2.5D sensor networks (Figure 7 (a)-(b)), in which each cross section in Z direction is same. This type of structure is stable and relatively not that challenge to print. Comparing to 2.5D structure, truly 3D structure is more difficult to print, however, we still successfully printed some samples, Figure 7 (c) shows a turbine impeller printed. Upon close inspection, each blade is clearly formed. Figure 7 (d) is a 3D lattice structure, this part needs support as shown in the process.

Upon close inspection, each blade is clearly formed. Figure 7 (d) is a 3D lattice structure, this part needs support as shown in the process. The challenges currently in our project lies in the frosting problem, which is mainly because of the temperature gradient around the print area. Upon investing to get a cold chamber, we expect this problem can readily be resolved.1617Figure 7 3D printed graphene oxide ice structures (a)-(b) Strain sensor network (c)-(d) Turbine rotor and truss 3D structure6.2 Aerogel formationGraphene aerogel is the terminal product to fabricate in this project. After the printing is finished, the prints are transferred to a -80 oC freezer through liquid nitrogen box. The samples are then put in a freeze drying machine for super critical drying. Figure 8 (a)-(b) shows some of the freeze dried GO samples. We found that constant chamber temperature is a key factor for successful samples.

Figure 8 (a)-(b) shows some of the freeze dried GO samples. We found that constant chamber temperature is a key factor for successful samples. To verify the success of formed aerogel, some performance tests are implemented, in Figure 8 (c), the aerogel is put on top of catkin, we can see the aerogel well stand without any depress on the catkin. In Figure 8 (d), the printed aerogel is connected between LED and power supply, the LED is lighted up. More properties test can be found in our previous work (Zhang et al., 2016).Figure 8 GO aerogels (a) 2.5D grid with different strut thickness (b) 3D GO aerogel periodic lattice of different size (c) Lightness test (d) Electrical conductivity test16187 ConclusionA novel directional freezing based 3D printing of graphene aerogel process has been presented.

The proposed system utilizes micro-dispensing technology to dispense GO water solution in a drop-on-demand manner onto a cold plate with specific temperature gradient, the dropped GOs freezes instantly upon landing the plate. By fully taking advantage of both 3D printing technology and freeze casting mechanism, macro and micro controllable truly 3D graphene aerogel can be fabricated. A process plan which covers the complete process and printing process for the developed system is proposed. An illustration of the directional freezing mechanism specifically for graphene oxide material is provided. A collection of successful samples are printed including 2.5D grid structure, special network, 3D graphene lattice, spatially varying 3D structure, and multiple material prints.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference4D printing method based on the composites with embedded continuous fibersQingrui Wang, Xiaoyong Tian, Lan HuangState Key Laboratory for Manufacturing Systems EngineeringXi’an Jiaotong UniversityXi’an, Shaanxi, 710049, ChinaAbstractMost of the current 4D printing technologies have the followingdefects: 1) the deformation shape is simple; 2) the deforming precision ispoor; 3) the deformation process is always uncontinuous. In this study, anew 4D printing process based on the composites with embeddedcontinuous fibers is proposed. In this process, a bilayer structureconsisting of the top layer of continuous fibers and the bottom layer withresin is 3D printed.

In this process, a bilayer structureconsisting of the top layer of continuous fibers and the bottom layer withresin is 3D printed. Due to the different thermal expansion coefficient andelastic modulus of the top and bottom layers, the structure will producebending deformation when the temperature changes. It is found that thecurvature value and the curvature direction of the composite structure canbe precisely controlled by the angle of the intersecting fibers. Theinfluence of fiber trajectory on curvature is studied, and then, thecontrollable deformation of any developable surface is achieved.1073Keywords: 4D printing, continuous fiber, composites, programmablemorphing1. Introduction4D printing is an additive manufacturing technology based on thedevelopment of 3D printing.

Introduction4D printing is an additive manufacturing technology based on thedevelopment of 3D printing. 4D printed material can produce controllabledeformation, and the factors that stimulate this deformation can betemperature, humidity and so on.[1] At present, there are three mainmethods for 4D printing. Shape memory materials are used as materialsfor 4D printing because they can change from temporary shapes to finalshapes at the transition temperature.[2-4] However, this deformation isoften discontinuous. Inspired by the deformation of pine cones and wheatawn, a new 4D printing method has been put forward.[5, 6] The rigidgranular materials, such as cellulose fibers, are embedded in a flexiblematrix, such as hydrogels. Because the rigid materials are inhomogeneous,the material will bend when it is placed in the water. However, thismaterial can only achieve simple and imprecise deformation.

Because the rigid materials are inhomogeneous,the material will bend when it is placed in the water. However, thismaterial can only achieve simple and imprecise deformation. If a variableexternal magnetic field or rheological field is applied during the printingprocess, a controllable orientation distribution of embedded particles canbe obtained, and the complex shapes can be obtained.[7, 8] There are stillseveral problems in this technology: first, the deformation accuracy of thematerial is not high; second, the magnetic field line can not be arbitrarilydistributed; third, the printing process is complex.1074In order to solve these problems, we developed the 4D printingtechnology based on continuous fiber embedded composites. Theorientation of continuous fibers is easier to control, which will lead tohigher deformation accuracy. In addition, this process is simpler and doesnot require complex outfield during the printing process.

In addition, this process is simpler and doesnot require complex outfield during the printing process. Throughtheoretical research and experimental verification, we have realized thedeformation of complex developable surfaces.2. Experimental section2.1 Materials and equipmentIn this paper, the flexible matrix material used is Polyamide66 (PA66,SunDcreate Corp. in China), and the continuous fiber material is carbonfiber (Tangu Corp. in China). The 3D printer used in this experiment isused to print continuous fiber reinforced composites.[9] The fiber andresin can be fed into the heating head at the same time and fullyimpregnated. The composite material is directly extruded from the nozzleoutlet. With this device, we can print out composite materials which canproduce controllable deformation.2.2 Preparation of continuous fibers embedded compositesThe composite structure in this experiment is a bilayer.

The bottomlayer is resin, and the top layer is a mixture of continuous fibers and resin1075(shown in Figure 1). Because these two layers have different elasticmoduli and coefficients of thermal expansion (CTEs), the compositestructure can produce bending deformation when the temperature changes.As the specimen just removed from the substrate has internal stress, theheat treatment is necessary. The heat treatment temperature is 170-210 °Cand the heat treatment time is 1-3 minutes.Figure 1 4D printing process of continuous fibers embedded Composites2.3 Curvature measurement of surfacesAccording to the experiment, when the fibers are distributed parallelon the surface of the composite, the material will be bent into acylindrical shape. Therefore, we can measure the degree of deformationby measuring the curvature of the cylinder.

Therefore, we can measure the degree of deformationby measuring the curvature of the cylinder. When the radian of thecylinder is less than π, the equation to solve the curvature is:kd2/=sin(kl)2/(1)where d is the chord length of the surface and l is the arc length of thesurface. When the radian of the cylinder is more than π, a caliper can be1076directly used to measure the diameter d and then to solve the curvatureby:k=2 d(2)3.

Results and discussion3.1 Deformation mechanism of the composite materialWhen the fiber bundles are distributed parallel on the resin surface,the composite structure will cause bending deformation due to thedifferent CTEs and elastic moduli of the continuous fibers and resin, asshown in Figure 2.Figure 2 The composite with parallel continuous fibers embeddedThrough the modulus of the resin is negligible compared to themodulus of the continuous fibers, the strain energy on the surface isapproximately equal to the sum of the strain energies of the two columns1077of fiber bundles:≈∆ e1 2E(2 ε ∆+∆ a2 ε b)(3)where Δε is the strain energy difference between when two columns offiber bundles exist together and when one column of fiber bundles existalone.For a point in the composite structure, assuming that its distancefrom the neutral surface is d, and the length of the neutral surface is l0,therefore:−′=∆ εεε 0=kd (−′lk ) 00(4)where k0 and ε0 are the curvature of the neutral surface and the strain atthe height of that point when a column of fiber bundles exists alone.

k’and ε’ are the curvature of the neutral surface and the strain at the heightof that point when two columns of fiber bundles exist together. Therefore,∝∆ e(ka−ka0)2+(kb−kb0)2(5)According to the Euler formula of differential geometry, thefollowing equation can be obtained:∝∆ e(ka−k 1cos2α )2+(kb−k 1cos2− αθ ())2(6)where θ is the angle between two fiber bundles and α is the anglebetween the direction of the principal curvature and one of the fiberbundles.1078For our 4D printing process, ka and kb are all the curvature when acolumn of fiber bundles exists alone, so ka=kb.

In order to make the △eminimum, the α and k1 (the principal curvature of the surface) can besolved: α      k 1==θ 2akcos2θ 2(7)To verify this equation, we printed rectangular composite panels withdifferent angles (30°, 60° and 90°) between two columns of fiber bundles.Figure 3 The direction of the curvature is along the angular bisector between fibersThe relationship between the size and the direction of the principal1079curvature and the angle of fiber bundles are shown in Figure 4 and 5.2.0Theoretical value Experimental value1.81.6ak /1k1.41.21.00306090θ(°)Figure 4 The relationship between the size of the curvature and the angle of fiber bundles40Theoretical value Experimental value) ° (1kf o n o i t c e r i d2000306090θ(°)Figure 5 The relationship between the direction of the curvature and the angle of fiber bundlesAccording to the theory and experiment, we can get the law that theprincipal curvature direction of the bilayer surface is in the direction of1080the bisector of the sharp angle between the fiber bundles, and theprincipal curvature value of the bilayer surface is proportional to sec2θ.With this rule, we can design more complex developable surfaces.3.2 Deformation design of a complex surfaceLet’s take a conical surfaceu=r (cosuv ,sinγuv cot , )as an example,where u, v are two parameters of the cone, and γ is the half vertex angleof the cone.

Solve the value and the direction of the non-zero principalcurvature, and the value of the non-zero principal curvature is:1 = kγcos u(8)It can be seen that the curvature of the position closer to the conicalvertex is larger, as shown in Figure 6a.

The direction of the curvature isalong the direction of the weft of the cone (Figure 6b).Figure 6 The size and the direction of the curvature of a coneThe cone is changed into a plane by isometric transformation, and1081the principal curvature can be expressed by the plane coordinate system(ρ, θ), as shown in Figure 7.1 =kγcot ρ(9)Figure 7 The principal curvature line after isometric transformationRearrange Equation 7 and 9, then the angle λ between the orientationof fibers and the direction of non-zero principal curvature can beobtained:cosλ±=k a k 1±=ρ k a γ cot(10)In polar coordinates,tanλ=′ ρ ρ(11)Rearrange Equation 10 and 11, then the equation for the fibertrajectory can be obtained:ρ d θ d±=ρ (cot akγ−ρ )(12)1082Solve the differential equation, therefore:ρ=2ka1 cotγ1[+sin(θ +±C)](13)where C is an arbitrary constant that determines the position of the curve,and the obtained fiber trajectories are shown in Figure 8.Figure 8 The obtained fiber trajectoriesIn order to verify this equation, we take the appropriate C values, sothat the distance between adjacent family of lines is near 8 mm.

Print thecomposite structure according to the designed carbon fiber trajectory andheat the material to get the deformed structures, as shown in Figure 9.1083Figure 9 The cone prepared by 4D printingIn order to verify the deformation accuracy, we printed many conicalsurfaces with designed fiber trajectories and cylindrical surfaces withparallel fiber trajectories. In each group of experiments, the printingparameters and the heat treatment parameters of the cone and the cylinderare exactly the same. The Ka calculated by the curvature of the bottomsurface of a conical surface is called the calculated curvature, and thecurvature measured by a cylindrical surface is called the experimentalcurvature. The two curvatures are compared, and the results are shown inTable 1.

The two curvatures are compared, and the results are shown inTable 1. It can be seen that the error between calculated curvature andexperimental curvature is between 3-8 %, indicating higher deformationaccuracy.Table 1 Comparison of experimental curvature and calculated curvatureGroup Experimental curvature Calculated curvatureError10.024770.022957.35%108420.015110.014573.57%30.066700.061198.26%40.048330.045436.00%According to this method, any developable surface can be preparedby this 4D printing process. The specific solution is still to get the maincurvature lines of the curved surface first, then transform it into a planecurve by isometric transformation, and finally get the fiber trajectory.4. ConclusionWe used 4D printing technology to prepare continuous fiberembedded composites.

ConclusionWe used 4D printing technology to prepare continuous fiberembedded composites. Due to the different elastic moduli and CTEs ofcontinuous fibers and resin, the composite structure will have differentflexural curvature at different temperatures. It is found that when the fiberbundles are aligned in parallel, the direction of principal curvature isalong the fiber orientation. When two columns of fiber bundles aredistributed on the surface of the resin, the direction of curvature is alongthe bisector of the angle θ between the intersecting fiber bundles, and thesize of the curvature is proportional to sec2θ. With this rule, we candesign more complex surfaces, such as a cone. This 4D printing processhas good designability and deformation accuracy.1085Reference[1] F. Momeni, S. M. Mehdi Hassani. N, X. Liu, J. Ni, MATER DESIGN 2017, 122, 42.[2] J. Wu, C. Yuan, Z. Ding, I. Michael, Y. Mao, SCI REP-UK 2016, 6, 1.[3] Z. Ding, C. Yuan, X. Peng, T.

 Printing of Soft Robotic Facial MusclesJiyu Cai*, Austin Vanhorn*, Casey Mullikin*, Jennifer Stabach*, Zach Alderman*, and Wenchao Zhou* The AM3 Lab, Department of Mechanical Engineering, University of Arkansas at Fayetteville Fayetteville, AR, United States of America; email: zhouw@uark.eduAbstract: 4D printing is an emerging technology that prints 3D structures with smart materials that can respond to external stimuli and change shape over time. 4D printing represents a major manufacturing paradigm shift from single-function static structures to dynamic structures with highly integrated functionalities. Direct printing of dynamic structures can provide great benefits (e.g., design freedom, reduced weight, volume, and cost) to a wide variety of applications, such as sensors and actuators, and robotics.

Soft robotics is a new direction of robotics in which hard and rigid components are replaced by soft and flexible materials to mimic actuation mechanisms in life, which are crucial for dealing with uncertain and dynamic tasks or environments. However, little research on direct printing of soft robotics has been reported. This paper presents a study on 4D printing of soft robotic facial muscles. Due to the short history of 4D printing, only a few smart materials have been successfully 4D printed, such as shape memory and thermo-responsive polymers, which have relatively small strains (~8%).

Due to the short history of 4D printing, only a few smart materials have been successfully 4D printed, such as shape memory and thermo-responsive polymers, which have relatively small strains (~8%). In order to produce the large motion needed for facial muscles, dielectric elastomer actuators (DEAs), operating like a capacitor with a sheet of elastomer sandwiched by two compliant electrodes and known as artificial muscle for its high elastic energy density and capability of producing large strains (~200%) compared to other smart materials, is chosen as the actuator for our robotic facial muscles. In this paper, we report the first fully 4D printed soft robotic face using DEAs. A literature review on DEAs is first presented. In order to select the right material for our soft robotic face, the performance of different silicone- based candidate materials is tested and compared.

In order to select the right material for our soft robotic face, the performance of different silicone- based candidate materials is tested and compared. A soft robotic face is then designed and fabricated using the selected material to achieve facial emotions by the motion of its lip and pupils actuated by the DEAs. This study demonstrates a 4D printed soft robotic face for the first time and the potential of 4D printing of soft robotics.1. IntroductionRobots are typically pictured as machines with rigid bodies and rigid motions. Soft robotics is a new direction of robotics in which hard and rigid components are replaced by soft and compliant materials to mimic actuation mechanisms in life. These soft robots are embedded with various sensors and actuators.

These soft robots are embedded with various sensors and actuators. One of the most important actuators in soft robotics is dielectric elastomer actuator (DEA), which operates like a capacitor with a sheet of elastomer sandwiched by two compliant electrodes, generating deformation under electric field. DEA is typically used as “artificial muscle” to mimic mammalian muscles due to its large strain and energy density. Natural muscle uses many complex mechanisms involving protein motion and regeneration from neural signal, while artificial muscles change in shape when an external stimulus is applied [1]. Some soft biomimetic robotics associated with artificial muscle are earthworm robots [2, 3], hexapod inspired robots [4, 5], Gecko climbing robot [6], and micro flying insect robot [7].As soft robotics develops over the past decades, the manufacturing methods are evolving from traditional machining, casting, and forging to more advanced approaches [8].

Although some soft1537robots were fabricated using traditional manufacturing processes, the demand for more advanced soft robotics has spurred the development of new manufacturing methods, such as shape deposition methods (SDM) and smart composite microstructures (SCM). SDM, a solid freeform fabrication (SFF) process, requires subtractive and additive manufacturing processes [9]. The subtractive manufacturing processes causes high energy consumption and large wastes of materials. SCM integrates the rigid structures with flexure joints and links, which is a fabrication process typically using laser micromachining methods [10]. These advanced manufacturing processes provide freedom in designing new types of 3D structured soft robots. Nonetheless, the manufacturing time and cost are still expensive for these methods. In addition, the complex manufacturing processes limit the complexity of robotic geometry and design freedom.

Nonetheless, the manufacturing time and cost are still expensive for these methods. In addition, the complex manufacturing processes limit the complexity of robotic geometry and design freedom. Therefore, in order to further improve the freedom of design, the ease of fabrication and the manufacturing efficiency in time and costs, simpler and more direct manufacturing technologies are desired for making soft robotics.3D printing technologies have made significant strides during the past decades enabling the direct fabrication of 3D structures. 4D printing, as an extension of 3D printing, is an emerging technology that prints 3D structures with smart materials, which can respond to external stimuli and change shape or properties over time. 4D printing represents a major manufacturing paradigm shift from single-function static structures to dynamic structures with highly integrated functionalities.

4D printing represents a major manufacturing paradigm shift from single-function static structures to dynamic structures with highly integrated functionalities. Direct printing of dynamic structures can provide great benefits (e.g., design freedom, reduced time and cost) for a wide variety of applications, such as sensors, actuators, and robotics. Due to its short history, only a few smart materials have been successfully 4D printed for various applications, including temperature-responsive polymers [11], shape memory polymers [12, 13], and ionic electroactive polymers [14]. And few publications are found on direct printing of dielectric elastomers for making soft robotics. Clearly more research is required in this promising direction due to the simple operation and high energy density of dielectric elastomers.

Clearly more research is required in this promising direction due to the simple operation and high energy density of dielectric elastomers. In this paper, we will demonstrate the promise of 4D printing of dielectric elastomers with the first fully printed soft robotic face that can smile upon application of voltage.This paper is organized as follows. In section 2, a comprehensive literature review is presented on dielectric elastomer actuators. In section 3, the detailed fabrication procedure of dielectric elastomer actuators is discussed, including material selection, fabrication process, and the issues encountered. In section 4, preliminary dielectric tests on simple dielectric elastomer actuators is performed to test their performance, the knowledge of which is then used for designing facial actuators. In section 5, a functional soft facial robot is designed and tested. Conclusions are given in section 6, and future works are discussed in section 7.2.

In section 5, a functional soft facial robot is designed and tested. Conclusions are given in section 6, and future works are discussed in section 7.2. Literature reviewDielectric elastomer (DE), a group of electroactive polymers (EAP), is a kind of smart material that can generate large strains under electric field. There have been many applications of DEAs since 1990s, such as robotics [15], active vibration control of structures [16], and energy harvesting [17], due to its lightweight and high elastic energy density. Two of the most common DE materials are acrylic and silicone elastomers. Both have their own advantages and disadvantages. Specifically, acrylic elastomers can produce relatively large strain up to 380% and have high dielectric constant, while silicones generally have high electromechanical response speed, long durability, and large range of operating temperature [18].

Most of silicone products are in liquid form or gel form, and can easily be turned into solid form via certain curing processes. Due to the1538ease of processing and customization, silicone has wider applications than acrylic elastomer, which is typically in solid form.The properties of dielectric elastomers play a significant role on their applicability for 4D printing applications. Some desirable properties for producing large strain include high dielectric constant, high breakdown strength, and low Young’s modulus. Many investigations have been reported on various aspects of DEAs, such as the geometry of electrodes [16], the breakdown strength [19-21], the effect of pre-strain on dielectric elastomers [22-24], and the enhancement of silicone dielectric and mechanical properties[1, 25-30].

One study demonstrates a heart shape of DEA, which could expand and remain in heart shape with fast response at a high frequency of 20 kHz [16].DEAs typically break down in three different ways, the electrical, the partial discharge, and the thermally initiated breakdown [19]. The partial discharge and thermally initiated breakdowns during the test period can be mitigated [19]. The electrical breakdown, also called dielectric breakdown, limits the maximum voltage that can be applied on DEAs, and thus limits the DEAs strain generation. Dielectric breakdown strength is also dependent on the material physical properties, such as mechanical stiffness, thickness, and pre-strain. The empirical results showed that the dielectric breakdown strength of silicone elastomer films as actuators can be enhanced by decreasing thickness or increasing pre-strain [20].

The empirical results showed that the dielectric breakdown strength of silicone elastomer films as actuators can be enhanced by decreasing thickness or increasing pre-strain [20]. In addition, un-cured or not fully cured silicone gel has a comparatively lower dielectric breakdown strength than fully cured silicone [21].The silicone elastomer generally has a maximum uniaxial elongation larger than 75%, and the pre-strain of elastomers have been shown by experiments and simulations as an effective approach to increase their maximum elongation [22]. The maximum strain of a uniaxial actuator made with Sylgard 186 silicone elastomer increased by 10 times from 7.7% to 80% when a 75% pre-strain was applied [24]. However, it was reported that excessive pre-strain could stiffen the elastomer and increase the requirement for actuation voltage. From experiments, Dow Corning HS3 RTV silicone DEA was shown to have a peak of elongation with a pre-strain of about 15%.

From experiments, Dow Corning HS3 RTV silicone DEA was shown to have a peak of elongation with a pre-strain of about 15%. In addition, from experiments, the softest elastomer is not necessarily the best choice to achieve the largest deformation, because the dielectric constant of the softest elastomer is not the highest [24].It is desirable to have a large dielectric constant for large electrostatic pressure, which can produce large mechanical strain. Many different approaches have been developed for increasing the dielectric constant of dielectric elastomers. Three approaches were explored, including random composites, field-structured composites, and new synthetic polymers [1]. To illustrate, random composite approach is to fill and disperse either solid (powders) or liquid materials into dielectric elastomer, such as ferro/piezoelectric ceramics, metal conductive particles, and organic polymers.

Field-structured composite approach is by mixing dielectric elastomers with some ferro/piezoelectric ceramics and curing them at an external electric field to form the dipolar molecular structure so as to increase dielectric constant. The polymer synthesis approach creates new molecular-structure for dielectric elastomer materials by copolymerization.For silicone elastomer, random composite is a typical approach for enhancing dielectric constant by studying the impact of the types and content of fillers on the DEAs performance. For example, some high dielectric constant materials were added in pure silicone to enhance its1539dielectric constant, such as ferro/piezoelectric ceramics (titanium dioxide and BaTiO3 powders) [25, 26] and multi-walled carbon nanotubes [27]. However, adding high dielectric constant materials may increase the elastic modulus and the decrease of the breakdown strength, which are not desirable for DEAs [26].

However, adding high dielectric constant materials may increase the elastic modulus and the decrease of the breakdown strength, which are not desirable for DEAs [26]. Softening the materials is another type of random composite method through adding plasticizers to enable large deformation at relatively low electric field [28]. In addition, there are examples demonstrating the synthesis approaches to improve silicone dielectric constant, such as copolymerizing copper-catalysed azide-alkyne 1,3-dipolar cycloaddition (CuAAC) [29] and polar cyanopropyl group polymers [30].Although significant progress has been made on improving the properties of dielectric elastomers, little research has been reported on direct printing of them. One study on 3D printing methods for DEAs was reported in 2008 at the University of California at Berkeley presented the 3D printing method for DEAs fabrication [31].

One study on 3D printing methods for DEAs was reported in 2008 at the University of California at Berkeley presented the 3D printing method for DEAs fabrication [31]. This study covered DEAs design, dielectric breakdown strength study and the design of printable pre-strain structures. Although their DEAs and printable pre-strain structures were designed for silicone elastomers, the dielectric test of silicone DEAs didn’t show any good results on mechanical strain or deformation. In addition, they did not demonstrate any soft robotic applications with their printed DEAs. In 2009, a contributive paper published in Japan presented a new approach to the fabrication of soft dielectric elastomer actuators using a 3D printing process [32]. They emphasized that the advantageous aspect of 3D printing processes is the capability to form complex three-dimensional structures, like actuators and sensors, through relatively direct fabrication methods.

A design of the pre-strain of 3D printed DEAs was also demonstrated.The literature survey reveals that extensive research has been conducted on the properties of DEAs (e.g., dielectric breakdown strength, effects of pre-strain, dielectric constant) while their fabrication methods have been largely overlooked. The very few research on the 3D printed DEAs has been very preliminary, and no soft robotics have been made with 3D printed DEAs. In addition, the design freedom endowed by 3D printing techniques is not well exploited. This paper is to demonstrate the first fully printed soft robotic face made with 3D printed DEAs.3. Dielectric Elastomer Actuators FabricationA typical DEA consists of two components: a dielectric elastomer film and compliant electrodes. For most of dielectric elastomers to produce a large strain, a common strategy is to pre- stretch the dielectric elastomer film to increase its dielectric constant.

For most of dielectric elastomers to produce a large strain, a common strategy is to pre- stretch the dielectric elastomer film to increase its dielectric constant. As a result, a rigid frame is typically required for pre-stretching the dielectric elastomer film. One typical configuration for building a pre-stretched DEA is shown in Figure 1. A dielectric elastomer film is stretched over a rigid circular frame with the electrodes area being masked off. Then conductive material is applied by painting or brushing over the film to make electrodes on both sides of the dielectric elastomer, which is sandwiched by the two electrodes. The novel part of our research is to construct the entire DEA configuration via 3D printing techniques, which provide the design freedom for the actuator’s geometry, in order to achieve the desired deformation of facial emotions.

The rigid frame is printed on a Fused Deposition Modeling (FDM) printer with ABS and the dielectric film and electrodes are printed on a Fab@Home printer, which is a multi-material 3D printer via dual syringe extrusion. The parts are manually assembled to apply pre-strain on the dielectric elastomer film.1540Figure 1. A schematic circular dielectric elastomer actuatorThrough applying voltage on the DE actuators, the electrostatic force between the electrodes squeezes the elastomer. The equivalent electromechanical pressure is the Maxwell electrostatic pressure, which is given by:𝑝𝑒𝑞 = 𝜀0𝜀𝑟𝑈2 𝑧2where 𝜀0 is the vacuum permittivity, 𝜀𝑟 is the dielectric constant of the elastomer, U is the applied voltage between electrodes and z is the thickness of elastomer membrane [33]. The primary challenge for achieving the apparent DEAs deformation of facial emotions is to select the right dielectric elastomer materials.

The primary challenge for achieving the apparent DEAs deformation of facial emotions is to select the right dielectric elastomer materials. Certain properties are desired based on the physical principles and our literature survey, such as high dielectric constant and dielectric breakdown strength for generating large pressure, and low Young’s modulus for obtaining large strain, as well as low viscosity for the ease of process. Silicone elastomer is chosen for its ease of processing. However, there are hundreds of silicone elastomer products on the market with various properties for different applications. Thus, it is necessary to conduct performance tests to select appropriate silicone elastomers that can generate large strain and operate at relatively low voltage for this facial robot.a. Dielectric Elastomer Actuator Configurations and MaterialsA simple DEA is designed for testing the performance of selected silicone elastomers.

Dielectric Elastomer Actuator Configurations and MaterialsA simple DEA is designed for testing the performance of selected silicone elastomers. The configuration of the test DEA is shown Figure 2, including silicone dielectric elastomer films, carbon conductive grease as electrodes, and circular rigid frames. All components of the DEA are 3D printed. Copper tapes are used as electrical wires to connect the DEA to a voltage source. Four silicone elastomers from different manufacturers are selected based on the aforementioned requirements, including low viscosity, high dielectric constant, high dielectric breakdown strength, and low Young’s modulus. These silicone elastomers’ specifications are shown in Table 1, and the elongation of materials can represent the capability of stretching. KE-1283, KE-3417 and KE-3494 silicone elastomers are all developed from ShinEtsu Chemical Company in Japan, and Sylgard 170 silicone elastomer is manufactured from Dow Corning Corporation.

KE-1283, KE-3417 and KE-3494 silicone elastomers are all developed from ShinEtsu Chemical Company in Japan, and Sylgard 170 silicone elastomer is manufactured from Dow Corning Corporation. These silicone elastomer films were printed on Fab@Home printer, which is an extrusion-based multi-material 3D printer.1541(1)Figure 2. A complete configuration of a circular DEATable 1. Specifications of silicone elastomersViscosity (cP)Elongation (%)Dielectric ConstantBreakdown Strength (kV/mm)KE-1283 KE-3417 KE-3494 Sylgard 1701950 5000 5000 2135300 260 250 -4 3.2 3.5 2.5425 19.7 25 18These circular rigid frame kits including clamps were printed on Ultimaker 2 with ABS plastic filament. In order to fit in the build area of Fab@Home printer (230x128x100 mm3), two circular frames with a diameter of 100 mm and height of 6 mm were designed, as shown in Figure 3.

In order to fit in the build area of Fab@Home printer (230x128x100 mm3), two circular frames with a diameter of 100 mm and height of 6 mm were designed, as shown in Figure 3. The idea of this frame design is to protect the printed electrodes and wire traces on both sides of dielectric elastomer film from touching the stage of the Fab@Home printer during printing process. The DE film is sandwiched between two frames, and three clamps can tightly hold the two frames together.Figure 3. SolidWorks model of assembled frames and clamps.Carbon conductive grease is used as the electrode materials, and printed on the dielectric elastomer film using the Fab@Home printer. For electrodes, it is critical to select a thin, extremely low-modulus electrode that can provide uniform charge distribution over the surface. MG carbon conductive grease was employed.b. Fabrication ProcessA Fab@Home multi-material printer is used for printing the silicone film and the electrodes.

MG carbon conductive grease was employed.b. Fabrication ProcessA Fab@Home multi-material printer is used for printing the silicone film and the electrodes. The adjustable process parameters are deposition rate, path speed, path width, path height, and syringe tip nozzle size, as explained in the Table 2. Uniformity and smoothness of printed dielectric elastomer films are dependent on these process parameters. The main procedures in the fabrication1542process are formulated, including material preparation, printing process, and curing post-process, as described below.Table 2.

The main procedures in the fabrication1542process are formulated, including material preparation, printing process, and curing post-process, as described below.Table 2. The description for processing parametersDeposition rateThe extruding rate (also extruding pressure)Path speedThe traveling speed of the print head while printingPath widthThe spacing between two parallel pathsPath heightThe single layer thicknessSyringe tip sizeTip nozzle size, also the diameter of extruded material linea) Material preparationBefore the printing parameters are determined, the materials have to be prepared for printing. One critical issue reported in literature is that air bubbles are easily trapped in viscous materials, inducing the inconsistence in material appearances and properties [26]. An example of a printed silicone film without sufficient degassing shows the bad quality of the print in Figure 4 below.

An example of a printed silicone film without sufficient degassing shows the bad quality of the print in Figure 4 below. To eliminate the air bubble issue, a vacuum chamber that can extract about 28.5 inches mercury pressure is used to degas the silicone or carbon grease.The results of the degassing process are influenced by two parameters: extracting pressure and duration. The maximum extracting pressure is determined by the power of pump, while only duration is a controllable parameter. Our experiments show 30 minutes of degassing is sufficient, and it is only effective for materials with viscosity of below 7000 cP, which is good for low and medium viscosity silicones and carbon conductive grease. Through observation, air bubbles in lower viscosity silicones are much easier to remove and come out much faster, such as KE-1283 and Sylgard 170 silicone elastomers.

Through observation, air bubbles in lower viscosity silicones are much easier to remove and come out much faster, such as KE-1283 and Sylgard 170 silicone elastomers. Another issue involved with degassing process is that fast- cure silicone may cure during the degassing process due to the rapid moisture loss in vacuum condition. Of the four selected silicone products, KE-3417 and KE-3494 elastomers cure fast, typically in one hour, while KE-1283 and Sylgard 170 elastomers cure in more than 5 hours. The alternative operation for fast-cure silicone is to degas in three iterations of 10 minutes, instead of one iteration of 30 minutes, and at each break the cured and solidified silicone must be removed so as not to block air bubbles coming out. In order to avoid this issue, the slow-cure silicone elastomers are preferred for easier processing, such as KE-1283 and Sylgard 170 silicones.Figure 4.

In order to avoid this issue, the slow-cure silicone elastomers are preferred for easier processing, such as KE-1283 and Sylgard 170 silicones.Figure 4. A printed silicone film with insufficient degassing1543b) Printing processThe printing process for silicone elastomer was conducted on the Fab@Home printer using a dual-syringe kit with 10 cc syringe barrels and 0.4 mm nozzle size tips, shown in Figure 5. The aforementioned processing parameters, such as deposition rate, path speed, path width, path height, and tip nozzle size, play an important role in determining the uniformity and smoothness of the printed film.abFigure 5. a) A syringe kit of 10 cc barrel and 0.4 mm nozzle tip. b) Dual syringe kit with two different silicone elastomers in barrelsThe impact of each parameter on the properties of print is explained below. Path height is associated with the film thickness, which determines how thick each layer will be and how many layers will be built up.

Path height is associated with the film thickness, which determines how thick each layer will be and how many layers will be built up. For this project, a single layer of elastomer with a various thickness from 100 µm to 500 µm is printed. Single layer can avoid potential air bubble issue between layers, and a thin film is desired to produce large electrostatic pressure according to Equation (1).Path width and tip nozzle size need to match up in order to print uniform and smooth films without any ridges or valleys, but not necessarily equal. For example, when a 0.4 mm nozzle tip is used, the path width needs to be around 0.4 mm. However, the material viscosity also needs to be taken into account. For low viscosity materials (~2000 cP), such as KE-1283 and Sylgard 170 silicone elastomers, the path width is recommended to be 0.5 mm because they are easy to spread out on the surface.

For low viscosity materials (~2000 cP), such as KE-1283 and Sylgard 170 silicone elastomers, the path width is recommended to be 0.5 mm because they are easy to spread out on the surface. For high viscosity materials (~5000 cP), such as KE-3417 and KE-3494 silicone elastomers, the path width needs to be 0.3 mm based on our experience.Deposition rate determines how much material is deposited on the substrate per second, while path speed means how fast the dual-syringe print head travels. They are related in some ways. The path speed is held at 30 mm/s, while the deposition rate varies from 0.4 mm3/s for low viscosity silicones to 1.6 mm3/s for high viscosity silicones. The reason for that is because higher viscosity silicones require higher extruding pressure, which is typically obtained by specifying a higher deposition rate for the Fab@Home machine.In addition, the levelness of platform is also a factor influencing the film uniformity, especially for thin films (100 µm).

Determining the parameters above is an experimental process to make the appropriate adjustment for different materials. Once the adjustment of parameters above are completed, a uniform and smooth film is obtained, as shown in Figure 6. Because it takes time for pressure to ramp up at the beginning of the print and to wind down at the end of the print, the print1544quality at the beginning and the end of the printing process is usually not as good. To solve this issue, two extra edges were designed for the silicone as shown in Figure 6.Figure 6. A good quality silicone elastomer film with extra areas.c) Post-processingFor curing, different silicone elastomers require totally different processes and conditions, shown in Table 3. The curing conditions may deviate from the conditions listed in Table 3, depending on the environment, such as the temperature and humidity.

The curing conditions may deviate from the conditions listed in Table 3, depending on the environment, such as the temperature and humidity. For example, KE-1283 silicone elastomer films should be cured in 2 hours at 50% relative humidity and 80 °C, but in high humidity circumstance (about 80% relative humidity in our lab), it took almost 4 hours to fully cure them. In addition to curing, pre-strain is necessary to obtain larger strain. Based on literature, higher pre-strain may produce larger strain, but the pre-strain on the silicone dielectric elastomers needs to be much less than the maximum allowed strain for the silicone film to remain soft. In this simple DEAs fabrication, 20% and 30% pre-strains for different silicone elastomers were successfully made, and their maximum allowed strains are more than 200%.

In this simple DEAs fabrication, 20% and 30% pre-strains for different silicone elastomers were successfully made, and their maximum allowed strains are more than 200%. The pre-stretched films need to be well aligned for precisely printing symmetric carbon grease electrodes on both sides of films, and then the round corner copper tapes are attached to the electrodes for connecting external voltage source. The round corner copper tapes can prevent the films from concentrated charging that may cause a local dielectric breakdown.Table 3. Curing conditions for different silicone elastomer KE-3417 Moisture 20 °C 2 hours *Note: Humidity in our lab is about 80% relative humidity (RH).KE-1283 Heat 80 °C 4 hoursKE-3494 Moisture 20 °C 6 hoursCuring Type Condition DurationSylgard 170 Heat 100 °C 20 minutes4.

Dielectric Elastomer Performance TestsThe dielectric elastomer performance tests were performed on a custom test setup with a high- voltage DC power supply developed by APEI Company at the University of Arkansas. The power supply was remotely controlled by LabVIEW to step up voltage at 200 V per 250 ms and to collect leakage current at each step. The highest voltage that the power supply can go is about 60 kV. Dielectric breakdown voltages were automatically detected and recorded, and the deformation of DEAs were manually filmed by a camera.1545After a silicone elastomer actuator was fabricated, a performance test was performed to measure the breakdown voltages and the maximum strains in order to compare with the performance of DEAs made with other silicones. Figure 7 shows an example of a printed DE film with KE-3494 silicone elastomer, which was pre-stretched by 1.2 times from a 55.8 mm diameter to a 67 mm diameter.

Figure 7 shows an example of a printed DE film with KE-3494 silicone elastomer, which was pre-stretched by 1.2 times from a 55.8 mm diameter to a 67 mm diameter. Black carbon conductive grease was printed as the electrodes on both sides.Figure 7. KE-3494 silicone elastomer actuator with 20% pre-strainDielectric breakdown strength measurement is an effective method to test the quality of printed silicone elastomer films and show the effects of pre-strain. Specifically, bad quality of silicone elastomer films caused by air bubbles or other issues commonly weaken their material properties, such as dielectric breakdown strength, dielectric constant, and Young’s modulus. In order to test the dielectric breakdown strength, both dielectric breakdown voltage and film thickness were measured. Table 4 shows the average of four thickness measurements for each silicone film by a micrometer and their tested breakdown voltage, as well as the tested and standard breakdown strength.

Table 4 shows the average of four thickness measurements for each silicone film by a micrometer and their tested breakdown voltage, as well as the tested and standard breakdown strength. The standard breakdown strength is tested by the suppliers for the silicone elastomers without pre-strain. In order to test the quality of printed film, the tested dielectric breakdown strength of silicone DEA without pre-strain was measured and compared with the standard breakdown strength as shown in the first row of Table 4. The tested dielectric breakdown strength is approximately the same with standard breakdown strength, which shows the DE film made with 3D printing techniques with degassing process has comparable quality to that made with traditional manufacturing processes. In addition, Table 4 indicates that pre-strain can positively increase the breakdown strength for each silicone elastomer.

1.2x pre-stretched Sylgard 170 silicone DEAs show 1.4% maximum strain and slightly normal dimension deformationb). 1.2x pre-stretched KE-1283 silicone DEAs show 5.9% maximum strainInitialEndingInitialEndingc). 1.2x pre-stretched KE-3417 silicone DEAs show about 2.1% maximum straind). 1.2x pre-stretched KE-3494 silicone DEAs show about 1.8% strain and obvious normal deformationInitialEndingse). 1.3x pre-stretched KE-3494 silicone DEAs show about 2.4% maximum strain and slightly normal deformationFigure 8. Performance tests for different silicone-based DEAsAnother important performance parameter we hope to obtain through the dielectric tests is the maximum strain for each silicone elastomer, which is crucial for making our facial robots. Figure 8 above shows the initial status and the status before breaking down for each silicone elastomer, as well as its measured maximum strain.

Figure 8 above shows the initial status and the status before breaking down for each silicone elastomer, as well as its measured maximum strain. All pre-stretched silicone DEA samples have the maximum strain less than 10%, which is similar with previously reported results by the conventional methods in literature. In addition, some of samples, such as Sylgard 170 and KE- 3494 silicone elastomers, showed normal (perpendicular to the planar film) deformation, which can be converted into planar deformation when higher pre-strain is applied. A comparison of different pre-strain rate on KE-3494 silicone DEAs indicates that larger pre-strain is helpful to generate more maximum strain and reduce the axial deformation. Among four silicone elastomers,1547KE-1283 silicone DEA shows the largest maximum strain (5.9%) at 20% pre-strain, and was selected as the dielectric elastomer film for the facial robot.5. Facial Robotic System a.

Facial Robotic System a. The idea of DEA-based facial robotic systemIn this paper, a facial robotic system was designed to obtain eye pupils expansion and lip curving activated by DEAs. The initial idea of design is that the expansion of pupils can be directly obtained from circular actuators which can expand isotropically and the smiley lip can be demonstrated from the relative deformation of points at different locations when an electric field is applied, shown in Figure 9. The lip curving is actuated to mimic real muscle. As shown in Figure 9, for circular lip actuators, points a and d locate at the top of the left and right actuators, and points b and c are below the middle actuator. The initial lip is a straight line connecting points a, b, c, and d. As voltage is supplied to system, the lip motion can be demonstrated by the deformation of these four points. Points a and d move upward while points b and c move downward to curve the lip to make a smiley face.

Points a and d move upward while points b and c move downward to curve the lip to make a smiley face. In this case, the actuators need to have large strain of at least 10% to demonstrate the facial emotions.a bc dFigure 9. A schematic facial deformation with circular pupil and lip actuatorsb. A modified facial robotic system design based on the DEAs performance testsAs shown in Figure 8, the maximum strain of the selected silicone dielectric elastomers are all less than 10% from the preliminary tests using simple DEAs. The largest strain (5.9%) was obtained from the KE-1283 silicone. Thus, facial robotic system is re-designed in attempt to maximize deformation with the relatively small strain of the DEAs. In order to demonstrate the facial emotion, a possible and feasible method is to increase the electrode size several times larger than in previous designs to obtain large deformation. Based on this idea, an alternative design was proposed as shown in Figure 13.

Based on this idea, an alternative design was proposed as shown in Figure 13. To clarify, the expansion of eye pupils are demonstrated by two circular DEAs at top, while the lip deformation is presented by the relative deformation between point 2 at the lowest edge of the large irregular-shaped electrode and two “static” points 1 and 3, shown in Figure 10-a. The deformation of electrodes illustrated by the dash line are calculated using the 5.9% strain of KE-1283 silicone elastomer. As electric field is applied, point 2 follows the bottom deformation of the large electrode moving downward, while points 1 and 3 that are far1548away from the electrode are expected to have negligible deformation. If the initial lip is designed to be the straight line connecting these three points, shown in Figure 10-b, the relative deformation of point 2 and points 1 and 3 would curve the lip downward to make a smiley face.

This design would also take advantage of the capability afforded by 3D printing to make arbitrary shapes.123a). Printed electrodes b). Facial features Figure 10. A modified design using irregular electrode for small strain actuatorsc. Feasibility test of facial robotic systemBesides increasing the electrode size to obtain large deformation, another possible method is increasing the pre-strain rate based on relevant literature and experiments. A 5.9% maximum strain was obtained on the KE-1283 silicone-based DEA with a 20% pre-strain, while the maximum allowed elongation of KE-1283 silicone is 300%. Thus, the capability of KE-1283 silicone under pre-strain condition has not been exploited well. In this facial robotic system, increasing pre-strain from previously 20 % to 50% was attempted to obtain large deformation.The performance for the facial robotic system was demonstrated via the dielectric test.

Dielectric breakdown strength and maximum strain measurements are shown in Table 5. The breakdown strength (35.3 kV/mm) of this 50% pre-stretched KE-1283 silicone elastomer film is slightly higher than the breakdown strength (34.9 kV/mm) of 20% pre-stretched sample in Table 4, and comparatively higher than the standard breakdown strength (18 kV/mm). The operation of the facial robotic system made with this DEA is shown in Figure 11 with the initial status and the final status before dielectric breakdown. However, due to the small strain of about 3.6%, the deformation cannot be clearly visualized. The obtained maximum strain about 3.6% is unexpected, and the potential reasons are only hypothesized. The electrostatic stresses of 20% and 50% pre- stretched KE-1283 silicone films are very close, since their dielectric breakdown strength are close to each other.

The electrostatic stresses of 20% and 50% pre- stretched KE-1283 silicone films are very close, since their dielectric breakdown strength are close to each other. Thus, it is more likely that small strain on the 50% pre-stretch KE-1283 silicone film is caused by the increased Young’s Modulus than that of 20% pre-stretch film. The weakness or defects may be amplified by the higher pre-strain (50%) than the 20% pre-strain, which could lead to worse film qualities (e.g., uniformity, smoothness) and reduce the performance of the DEAs.1549Another possibility is that higher pre-strain may make the silicone film stiffer so that cannot produce larger strain, which was reported in literature.Table 5. Dielectric test results for the facial DEAFilm Thickness (mm)Tested Breakdown Voltage (kV)Tested Breakdown strength (kV/mm)Standard Breakdown Strength (kV/mm)Maximum Strain (%)KE-1283 pre-strain 50%0.2157.635.3183.6InitialEndingFigure 11.

Performance of facial robotic system6. Conclusions We reported a fully 4D printed soft robotic face using DEAs in this paper for the first time. First, we provided a comprehensive literature survey on DEAs. Then we formulated a procedure for fabricating DEA using 3D printing techniques. It was shown that the 3D printed dielectric elastomer film has comparable quality to that made with traditional manufacturing processes. In order to select the right material for our soft robotic face, we tested the performance of the DEAs made with different commercial silicone elastomers. KE-1283 was shown to have the best performance among the tested materials. A soft robotic face was then designed and printed with DEAs to actuate the pupils and the lip for obtaining a smiley face. Upon application of voltage, a smiley face was successfully achieved.

A soft robotic face was then designed and printed with DEAs to actuate the pupils and the lip for obtaining a smiley face. Upon application of voltage, a smiley face was successfully achieved. In conclusion, this paper demonstrated a fully printed soft robotic face and the feasibility of using 3D printing for making DEAs with complex geometry and soft robotics.7. Future Work 4D printing is a burgeoning field. Base on this research, there is still a huge room to learn and improve the current design. Models need to be developed for understanding and predicting the performance of DEAs to assist the design of the DEAs and soft robots. Further investigations are also needed to optimize process parameters and the geometry of actuators to improve the performance of DEAs and enable more sophisticated soft robotic applications.15508.

Acknowledgement We gratefully acknowledge the financial support from the University of Arkansas, through the startup fund provided by the Vice Provost Office for Research and Economic Development. Any opinions, findings, and conclusions or recommendations expressed in this publication are those of the authors and do not necessarily reflect the views of the University of Arkansas. We also acknowledge assistance from all AM3 lab members, and appreciate APEI Company for providing testing equipment.References[1] F. Carpi, D. De Rossi, R. Kornbluh, R. E. Pelrine and P. Sommer-Larsen, Dielectric Elastomers as Electromechanical Transducers: Fundamentals, Materials, Devices, Models and Applications of an Emerging Electroactive Polymer Technology. Elsevier, 2011.[2] A. Menciassi, S. Gorini, G. Pernorio, L. Weiting, F. Valvo and P. Dario, "Design, fabrication and performances of a biomimetic robotic earthworm," in Robotics and Biomimetics, 2004. ROBIO 2004.

14, pp. 487-494, 2007.[22] Y. Liu, L. Liu, Z. Zhang and J. Leng, "Dielectric elastomer film actuators: characterization, experiment and analysis," Smart Mater. Struct., vol. 18, pp. 095024, 2009.[23] G. Yang, G. Yao, W. Ren, G. Akhras, J. P. Szabo and B. K. Mukherjee, "The strain response of silicone dielectric elastomer actuators," in Smart Structures and Materials, 2005, pp. 134-143.[24] S. Akbari, S. Rosset and H. R. Shea, "More than 10-fold increase in the actuation strain of silicone dielectric elastomer actuators by applying prestrain," in SPIE Smart Structures and Materials Nondestructive Evaluation and Health Monitoring, 2013, pp. 86871P-86871P-13.[25] F. Carpi and D. D. Rossi, "Improvement of electromechanical actuating performances of a silicone dielectric elastomer by dispersion of titanium dioxide powder," Dielectrics and Electrical Insulation, IEEE Transactions On, vol. 12, pp. 835-843, 2005.1552[26] Z. Zhang, L. Liu, J. Fan, K. Yu, Y. Liu, L. Shi and J.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed Paper5-AXIS SLICING METHODS FOR ADDITIVE MANUFACTURING PROCESSSajan Kapil*, Seema Negi*, Prathamesh Joshi*, Jitendra Sonwane*, Arun Sharma*, Ranjeet Bhagchandani* and K. P. Karunakaran*Mechanical Engineering Department, Indian Institute of Technology Bombay, Powai, Mumbai, 400076, Maharashtra, IndiaAbstractIn metallic Additive Manufacturing (AM) processes such as Hybrid Layered Manufacturing (HLM), it is difficult to remove the support material used for realizing the overhanging/undercut features. Multi-axis kinematics can be used to eliminate the requirement of the support mechanism. In this work, two slicing methods have been proposed which utilize the benefits of multi-axis kinematics to eliminate the support mechanism.

In this work, two slicing methods have been proposed which utilize the benefits of multi-axis kinematics to eliminate the support mechanism. In the first method, planar slicing is used and the overhanging/undercut features are realized while keeping the growth of the component in the conventional Z-direction. In the second method, non-planar slicing is used, and the growth of the component need not necessarily be in the Z-direction; it can also be conformal to the selected feature of the component. Both these methods are explained through a case study of manufacturing an impeller by the HLM process.Keywords: Hybrid Layered Manufacturing, Rapid Prototyping, 5-Axis Cladding, 5-Axis Slicing, Non-planar SlicingIntroductionHybrid Layered Manufacturing (HLM) uses cladding units such as Metal Inert Gas (MIG), Tungsten Inert Gas (TIG) and LASER for material addition and conventional machining operation for material subtraction.

This synergic combination of material addition and subtraction process makes HLM an optimal method for metal based Additive Manufacturing (AM) processes. In our indigenous HLM process, the cladding units are retrofitted with the existing 3/5-axis CNC machines without disturbing their other capabilities, see Figure 1.(a) 3-Axis HLM(b) 5-Axis HLMFigure 1: Experimental setup of Hybrid Layered Manufacturing process1886Figure 2: Flow chart of Hybrid Layered Manufacturing processA typical flow chart of the HLM process is shown in Figure 2. For the geometric realization of a component, a valid CAD model is used as input to the HLM software which generates NC files corresponding to each layer as output. The generated NC files are then used to move the cladding units in the required region on the substrate for the realization of each layer. After deposition of each layer, three in-situ processes are used viz; (1) Face-milling (2) Pneumatic hammering and (3) Induction preheating.

After deposition of each layer, three in-situ processes are used viz; (1) Face-milling (2) Pneumatic hammering and (3) Induction preheating. The Face-milling operation is used to maintain the Z-accuracy, remove the scallop and oxide scale. The pneumatic hammering is used for relieving the stresses developed during the solidification of deposited material. An induction-based preheating system is then used to heat up the prebuilt layer before deposition of next layer. By repeating these steps for each layer a near-net shape of the geometry can be realized. This near-net shape is then finished by milling to obtain the final finished product. The cladding units can be retrofitted with the 5-axis CNC machine to eliminate the need for support material, Figure 1b. By utilizing the multi-axis kinematics the relative position between cladding unit and a prebuilt layer is set in such a way that the overhanging/undercut features can be deposited without using any support mechanism.

Several strategies have been proposed1887for multi-axis AM by many research groups. A literature survey has been carried out which is summarized in the following paragraph.The AM processes which use multidirectional deposition can reduce the amount of sacrificial material deposited. A methodology has been proposed by Singh and Dutta [1] for 5- axis slicing algorithms of a component. The main objective of the work was to select the build direction of different subvolume such that the support material can be eliminated/reduced and a collision -free deposition tool path can be generated. A similar approach has been adopted by Sundaram and Choi [3] for 5-Axis laser aided Direct Metal Deposition (DMD) process. They build the component in the +Z direction and the orientation of the object changes for deposition of the different subvolumes. In the work of Zhang [2], an algorithm has been developed for adaptive slicing for a 5-axis hybrid Laser Aided Manufacturing (LAM) process.

In the work of Zhang [2], an algorithm has been developed for adaptive slicing for a 5-axis hybrid Laser Aided Manufacturing (LAM) process. The overhang features in a hybrid LAM process are produced without support material by using two strategies: (a) Deposition of overhang on transition wall, (b) Deposition of overhang based on surface tension. 5-axis Hybrid Layer Manufacturing systems, where the deposition is done along with a subtraction process are capable of producing non-uniform layers thickness [4-6]. Hence with the integration of multi-axis deposition and machining processes on the same workstation, a hybrid system is able to produce complicated geometry. In the work of Ren et al. [7], the process planning strategies for 5-axis deposition is proposed by following steps (a) spatial decomposition (b) slicing of the part (c) tool path generation. In the work of Singh and Dutta [8], a method has been proposed for slicing in non-planar/multi-direction called "Offset Slicing".

In the work of Singh and Dutta [8], a method has been proposed for slicing in non-planar/multi-direction called "Offset Slicing". As the multi-direction deposition allows the deposition of 3D slices, these three- dimensional slices should be parallel to each other and the base (i.e. offset) to maintain a uniform thickness. Ruan et al. [9] have proposed a method for deposition of non-parallel and non-uniform layers to reduce the staircase effect. To deposit such 3D layers, this method uses an empirical model to predict the layer thickness. In the work of Ren et al. [10], an integrated process planning for Hybrid manufacturing system is presented. A hybrid process planning software has been developed which has the capability of automatic alignment of layers to avoid a support structure, automatic collision detection with a pre-deposited layer, adaptive tool path generation to avoid the voids in a contour parallel toolpath and automatic 5-axis finishing machining after deposition.

As discussed earlier in [4] to produce non-uniform slicing, if the machining is done on an STL file the accuracy may not be good hence Zhang et al. [11], proposed a method for improving the accuracy of 5-axis machining of each layer. A Continuous 5-axis laser cladding process has been recently studied by Calleja et al. [12]. They have performed the experiments to study different toolpath strategies and to arrive at an optimal toolpath considering quantitative criteria such as the deposition rate, the wetting angle and the height and width clad tracks.From the above literature review, it can be summarized that in all the 5-axis based deposition methods a 5-axis machining operation is required to develop non-uniform layer. In this work, two unique slicing methods have been proposed for depositing the layers continuously through 5-axis.

In this work, two unique slicing methods have been proposed for depositing the layers continuously through 5-axis. HLM process has been used for the experiments and the validation of the developed slicing procedure.Methodology and IllustrationIn this work, the slicing procedures have been classified on the basis of the kinematics involved in the deposition of a layer viz; (a) 3-axis slicing and (b) 5-axis slicing procedure. Both the slicing methods are explained in the following sections18883-Axis SlicingIf only 3-axis kinematics is involved for the realization of a layer then it will be called 3-axis slicing method. To realize a layer, the slice is consist of two tool paths viz; (a) outer contour deposition and (b) inner area-filling. For each layer first, the outer contours are deposited then the inside area is filled. Figure 3 shows the steps involve in 3-axis slicing method.

For each layer first, the outer contours are deposited then the inside area is filled. Figure 3 shows the steps involve in 3-axis slicing method. For inside area-filling, several methods are available such as direction parallel, contour parallel and other area-filling strategies [13].(a) CAD model of the component(b) 3-axis slicing of a component(c) Intersecting outer contours in a slice(d) Inside area filling by contour parallel toolpathFigure 3: 3-axis slicing procedureWhile depositing the outer contours it should be noted that if the overhanging feature is beyond a limit (which depends upon the surface tension of the material), then it will be required an external support mechanism. For example, a slant wall of Mild-steel with an angle from vertical can be easily manufactured by 3-axis slicing without using any support of material. Adaptive Slicing is one method of manufacturing the overhanging features on 3-axis.

Adaptive Slicing is one method of manufacturing the overhanging features on 3-axis. In this method, the CAD model is sliced with varying layer thicknesses. The region of the object with a higher gradient with respect to the build direction is sliced with a smaller layer thickness and vice versa. If the current layer overhangs the previous one beyond a limit then adaptive layer thickness method can be utilized as shown in Figure 4.30°1889(a) Variable layer thickness in circular duct(b) Actual deposition of the ductFigure 4: Adaptive slicing method5-Axis SlicingAs discussed above, if the overhanging feature is beyond a limiting value then an external support mechanism will be required. To eliminate the need for the external support mechanism a 5-axis kinematics can also be utilized. In this section, two methods have been proposed for continuous 5-axis deposition viz; (a) planar slicing and (b) non-planar conformal slicing.

In this section, two methods have been proposed for continuous 5-axis deposition viz; (a) planar slicing and (b) non-planar conformal slicing. Both the methods are described below:Planar SlicingIn the case of planar slicing method, the component is still built in the similar way as discussed in the previous section for 3-axis slicing. The only difference will be in the deposition of the outer contours. While depositing the outer contours the cladding torch vector need to be aligned with the tangent vector as shown in Figure 5a. Once the outer contours are deposited by tilting the cladding torch or substrate along the tangent vector a bowl kind of shape is prepared, see Figure 5b. Now the inside area can be filled as earlier using the 3-axis kinematics, as shown in Figure 5c. Each layer is deposited by repeating the same steps.As discussed above the torch vectors need to be calculated for outer contour depositionin each slice.

Figure 6b shows the slicing of the cylindrical part and Figure 6c shows the outer intersecting contours in an intermediate slice. After the realization of inside cylindrical feature on 3-axis, the outer feature needs to be realized on its periphery. For that, the base (which is cylindrical in this case) will be offset by the required layer thickness as shown in Figure 6d. One can notice now that the each slice is a cylindrical surface with a different radius equal to the offset value. Figure 6e show an intermediate non-planar slice with the intersecting contours. Similar to the previous method the material can be deposited along these contours and then the inside area can be filled as shown in Figure 6f.All three described methods (3-axis adaptive slicing, 5-axis slicing, and non-planar conformal slicing) has been demonstrated as a case study by manufacturing an impeller.

The results obtained for this case study has been discussed in the next section.1891(a) CAD model of the component(b) Slicing of the inside cylindrical feature(c) An intermediate slice with intersecting contours(d) Slicing of the outer feature with non-planar cylindrical slices(e) An intermediate non- planar slice with the intersecting contours(f) Toolpath generation for the material deposition in the non- planar sliceFigure 6: Feature-based non-planar conformal slicing methodResults and DiscussionAn impeller has been manufactured using HLM process as a case study by the above discussed three methods. Figure 7a shows the CAD model of an impeller to be manufactured by using adaptive 3-axis slicing method. One can observe that in the build direction, on the top of the impeller there is a large overhanging feature. Hence an adaptive slicing will be required to keep the constant overhang.

One can observe that in the build direction, on the top of the impeller there is a large overhanging feature. Hence an adaptive slicing will be required to keep the constant overhang. The variable layer thickness from 2mm to 0.5mm was obtained by an in-situ face milling operation. Figure 7b shows the adaptive slicing of the impeller with the region of smaller layer thickness. Figure 7c show the intermediate stage of the impeller after deposition of few constant thickness layers (2mm). The final near-net shape obtained by the deposition is shown in Figure 7d. It was observed that if the layer thickness becomes very small (0.5 mm) then during the next layer deposition it starts melting and hence the material starts falling down. One can observe the rounded edges of the impeller in Figure 7d. This method could not produce a very accurate near-net shape of the impeller.

One can observe the rounded edges of the impeller in Figure 7d. This method could not produce a very accurate near-net shape of the impeller. Hence another method, 5-axis slicing was used which is described in the following paragraph.1892(a) CAD model of the scaled impeller(b) Adaptive slicing of the impeller(c) Deposition of impeller an intermediate slice(d) Final near-net shapeFigure 7: Realization of the impeller using 3-axis adaptive slicingFigure 8a shows the CAD model of the impeller to be manufactured by 5-axis planar slicing. As discussed earlier the planar slices are done as shown in Figure 8b. The intersection between each layer and component was obtained with required torch vector for the overhanging feature. First, the outer contours are deposited using 5-axis kinematics by tilting the base plate accordingly. Figure 8c shows the deposition of the outer contour in an intermediate slice. Then the inside area is filled using a 3-axis kinematics, see Figure 8d.

Figure 8c shows the deposition of the outer contour in an intermediate slice. Then the inside area is filled using a 3-axis kinematics, see Figure 8d. Repeating these steps for all the slices will result in a near-net shape of the impeller as shown in Figure 8e. Compare to the previously deposited near-net shape by 3-axis slicing method, this near-net shape is found to be more accurate. The near-net shape is then final milled and the finished component is shown in Figure 8f. The final product’s weight is 1.015 Kg while the weight of near-net shape is 1.826 Kg. The weight of the bounding cylinder of 170mm x 80mm is 4.901 Kg. Hence the material saved as compared to the purely subtractive process is 4.901 - 1.826 = 3.075 Kg.

The weight of the bounding cylinder of 170mm x 80mm is 4.901 Kg. Hence the material saved as compared to the purely subtractive process is 4.901 - 1.826 = 3.075 Kg. The time was also recorded throughout the deposition process and it was found that it took approximately 4hr to realize the near-net shape of the impeller.(a) CAD model of the impeller(b) Planar slicing of the impeller1893(c) Outer contour deposition of an intermediate slice(d) Area filling of an intermediate slice(e) Final Near-net shape(f) Final finished componentFigure 8: Realization of the impeller using 5-axis planar slicingTo further improve the near-net shape of the impeller obtained by the planar 5-axis slicing method, a non-planar conformal slicing method was used. As described earlier, first the features are identified based on their buildability in the different build direction.

As described earlier, first the features are identified based on their buildability in the different build direction. Figure 9a shows the decomposed CAD model of the same impeller into two features viz; (a) inside the hub and (b) blades. The inside hub can be manufactured using the 3-axis kinematics, see Figure 9b. After deposition of inside hub, its top surface was considered to be the first slice for the blade deposition. The hub surface is then offset by 1mm up to the radial height of the blades, Figure 9c shows all the non-planar slices for the realization of the blades on the hub. An intermediate slice with deposition toolpath and it actual deposition is shown in Figure 9d and 9e respectively. Depositing the blades on each offset non-planar slice will result in the near-net shape of the impeller shown in Figure 9f.

Depositing the blades on each offset non-planar slice will result in the near-net shape of the impeller shown in Figure 9f. One can clearly observe that this near-net shape is much closer to the CAD than the one created by the last method of 5-axis planar slicing.(a) Feature identification(b) Deposition of hub1894(c) Non-planar slicing for the blade deposition(d) Deposition path in an intermediate slice(e) Deposition the intermediate layer of blades(f) Near-net shape of the impellerFigure 9: Realization of the impeller using 5-axis non-planar slicingConclusionsFor Additive Manufacturing (AM) process three slicing methods have been described in this work, viz; 3-axis adaptive slicing method, 5-axis planar slicing method, and 5-axis non- planar slicing method. Each method has been explained with the help of a suitable example.

Each method has been explained with the help of a suitable example. An impeller was manufactured using Hybrid Layered Manufacturing (HLM) process, as a case study to demonstrate the capability of the proposed slicing methods. The near-net shape produced by 5-axis planar slicing method has the consistent deposition on each layer without any toolpath error. As compared to conventional machining it has been found that for this case study the saving in material is approximately 63%. The near-net shape produced by the 5-axis non-planar slicing method is found to be the best among all three methods. The developed slicing methods are not only limited to HLM process but can be used for other AM process such as SLS, FDM etc. The parameters that controls the adaptive slicing has to be investigated in future work.References1. P. Singh and D. Dutta, 2001, "Multi-Direction Slicing for Layered Manufacturing," Journal of Computing and Information Science in Engineering, 1(2), pp. 129 - 142.2.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA comparison of mechanical properties from natural and process-induced interfaces in filament extrusion AM of polymer blendsCamden A. Chathamand Donald W. Benza**Advanced Engineering Division, Savannah River National Laboratory, Aiken, SC USA E-mail: camden02.chatham@srnl.doe.govAbstractPolymer blends are commonly tuned for specific applications to achieve desired properties otherwise inaccessible or prohibitively expensive to obtain via homopolymers. The interfacial characteristics of the polymer A-polymer B interface and resultant domain sizes govern key performance properties. Micro- and meso-scale morphology forms through the interplay of surface forces between the polymers and between each polymer and the surrounding atmosphere.

Micro- and meso-scale morphology forms through the interplay of surface forces between the polymers and between each polymer and the surrounding atmosphere. Analogously, the layer-layer and road-road interfaces of material extrusion (MEX) additive manufacturing (AM) govern key performance properties of printed parts. This work explores the effect of layer height on the thermomechanical performance of polystyrene (PS)-polycarbonate (PC) blends. Filament is prepared from a 50/50 weight ratio of the two polymers and compared against dual-nozzle printing where every layer alternates between PS or PC homopolymer forming a part with an overall 50/50 polymer ratio. Typical indicators of polymer blend compatibility are also studied.IntroductionAdditive manufacturing’s (AM) layerwise manufacturing paradigm is commonly touted for its ability to fabricate complex geometries without shape-specific tooling.

This fabrication method enables lattices to be fabricated just as easily as solid blocks leading to the summary statement that “complexity is free” in AM. However, material behavior at interfaces is significantly different from that observed in the bulk. Therefore, interfacial bonding (both interlayer and intra- layer) is a critical aspect of AM processes to understand. Interfacial strength in composite and adhesive systems are well-studied due to the regularity at which part failure occurs at the interfaces between dissimilar materials in part due to the interface being weaker than regions where bulk behavior dominates.1 When using the fused filament fabrication (FFF) sub-technology of material extrusion (MEX) to fabricate parts from polymer blends, there exist two types of interfaces: (i) domain boundaries between constituent polymers (microscale) and (ii) process-induced boundaries between layers and roads (mesoscale).

These interfaces are hierarchical for polymers blended during filament production as the domain boundaries will be substantially smaller than the typical layer thickness of 100’s of microns.Several FFF printer manufacturers have developed machines with multiple extrusion nozzles enabling multiple materials to be incorporated into a single part, including the Ultimaker S32 used for the present work. Although other authors have reported on the interfacial strength,3–5 they fabricated specimens with a single dissimilar polymer layer interface. This method enables ease of measurement for interfacial strength; however, it provides no insight into whether a dual- extrusion part can approximate the properties of a polymer blend prepared in advance. To the authors’ knowledge, no published work compares the mechanical performance of a traditionally2blended filament with a structure of alternating homopolymer layers comprising an analogous polymer blend ratio.

Figure 1 depicts a cartoon representation of parts produced from traditionally blended filament and alternating homopolymer layers. This can be considered an expansion of nano/microlayered coextrusion6,7 as it promotes the macro-scale exhibiting of interfacial behavior and properties relative to bulk behavior and properties by decreasing the layer size and increasing the number of interfaces. Current FFF technologies use layer heights much larger than the 10s of nanometers usually needed for geometric confinement behavior,6 but the relative increase in volume of interphase polymer compared with bulk molding technologies is significant.Figure 1: Cartoon depiction of a generic part made from (a) traditionally blended filament and (b) alternating homopolymer layers via FFF.

Subfigure (c) depicts the alternating homopolymer layers method using thinner layers producing more interfaces.Self-segregated domains occur in polymer blends due to chemical differences between constituent polymers. These differences are often characterized using the Interaction Parameter (χ) with miscible blends having a negative value and compatible blends having values less than 0.5.8 The reported interaction parameter for PC and PS is 0.038 by Kim and Burns.9 Phase diagrams describing the resulting blend morphology from various combinations of polymers are the subject of painstaking scientific work. Such work can be leveraged for tailored self-assembly into morphologies desired for specific applications.

Such work can be leveraged for tailored self-assembly into morphologies desired for specific applications. Various morphologies can be achieved for the same polymer blend ratio through altering processing conditions.When blending two polymers, one might expect the performance behavior of the blend to be some combination of the performance from the two constituents and the newly created, third mixed phase. The study of the characteristic length scales and how those domains affect final part performance has been published by Chevallier, et al.,10 Tanaka, et al.,11 and Ayoub, et al.12 They report changes in observed domain sizes with changes to composition and processing.

Other researchers have found through transmission electron microscopy (TEM) that the width of the PS- PC interface is 32 nm.13 Often phase separated behavior shows a profound effect on melt viscosity, which in turn alters solid-state mechanical properties.11 Given the importance of these characteristic lengths of phase separation, the present work demonstrates the impact of nominally equivalent blends of polystyrene and polycarbonate by either blending the two components in a twin screw extruder prior to making filament or else extruding alternating layers of each homopolymer. Each method produces samples with distinct spatial arrangement of each blend3component. Traditionally mixed samples via twin-screw extrusion, like several of the literature- cited studies, are expected to have PS droplets dispersed inside a continuous PC phase14 or else a co-continuous phase;15 samples prepared from alternating layers are expected to largely retain their layered structure.

The difference in response to both oscillatory shear loading and creep recovery are explored and reported for these materials.ExperimentalMaterials, Filament Preparation, and FFF Printing.Polystyrene (Sigma Aldrich) and bisphenol-A polycarbonate (3DXTECH) were sourced in pellet-form for in-house blending and filament preparation. Homopolymer filament was prepared using a Filabot EX2 filament making system. 2.5 ± 0.1 mm filament was extruded at 245 °C. A 50/50 w/w PC/PS blends were prepared using a Liestritz twin screw extruder prior to re- pelletization. Filament was prepared from the blended pellets at 245 °C using the Filabot EX2 filament extruder. All samples were fabricated using an Ultimaker S3 dual extrusion FFF machine with a bed temperature of 110 °C, nozzle temperature of 260 °C, nozzle diameter of 0.4 mm, and print speed of 10 mm s−1.CharacterizationDynamic mechanical analysis (DMA).

A Mettler Toledo DMA/SDTA 1+ with shear clamping assembly was used to quantify viscoelastic properties of prepared blended polymers. Test specimen cylinders were prepared by FFF AM in the ZYX orientation (i.e., shear force applied parallel to layer interfaces). Test specimen geometries are presented in Table 1; all geometries afford a geometry factor within appropriate measurement limits according to the Mettler Toledo user manual.The Mettler Toledo DMA/SDTA 1+ operates by switching between stress and strain controlled measurement modes. Therefore, the user sets both a force amplitude and a displacement amplitude when creating a measurement method and the oscillation cycle ends whenever either criterion is met (i.e., a logical “OR”). The authors report both amplitudes for completeness as “XX N force amplitude or YY µm displacement amplitude” to emphasize the logical “OR” nature of the method.Layer Height [mm]Table 1: Designed dimensions of FFF printed samples.

Thickness [mm]Number of LayersDiameter [mm]Geometry Factor [m−1]0.30 0.1516 164.8 2.415 1013.581 15.2794Figure 2. Photographs depicting printed shear DMA samples. (a) Comparison between the differently sized samples to account for changing layer height at constant number of interfaces. (b) Comparison of 150 µm layer height samples among the four materials tested.An amplitude sweep was conducted at room temperature between 0.01 and 1.0 mm to determine the linear viscoelastic region. Specimens were first evaluated at room temperature by frequency sweep between 0.01 and 100 Hz at an amplitude of 2 µm before being heated at 3 °C min−1 at 1 Hz and 2 µm mm to 220 °C. Additionally, samples were tested for their creep response by applying a 30 N force offset for 6 h then resetting the offset to 0 N for 8 h. Creep recovery was evaluated at both 128 °C and 30 °C.Differential Scanning Calorimetry (DSC).

Creep recovery was evaluated at both 128 °C and 30 °C.Differential Scanning Calorimetry (DSC). A Mettler Toledo DSC 3+ was used to evaluate the thermal transitions of filament made in-house. Samples were heated at 10 °C min−1 between 0 °C and 250 °C for two heating and cooling cycles before being heated at 20 °C min−1 for a final heating cycle. Initial heating cycles were used to erase any thermal memory from the sample and the final heating cycle was used to determine Tg. The midpoint method was used for determination of glass transition temperature.5Results and DiscussionImpact on the glass transitionFigure 3 reports glass transition temperature obtained by both DSC and DMA. It is typical for Tg measured via DMA (via peak of the tan(δ) curve) to be approximately 10 °C above the value obtained via DSC (midpoint of the heat flow baseline shift).

It is typical for Tg measured via DMA (via peak of the tan(δ) curve) to be approximately 10 °C above the value obtained via DSC (midpoint of the heat flow baseline shift). The data collected from the PS and PC homopolymers follow the expected trend.Figure 3: (a) tan(δ) versus temperature for printed homopolymers and blends. (b) Second heat of DSC depicting Tg and melt behavior for homopolymer filament.The Fox equation (Eq 1) states that the expected Tg for a miscible polymer blend is the sum of the weighted inverse temperatures of the component homopolymers. In Equation 1, the a and b designators indicate the two homopolymers present in the blend; ω is the weight fraction of that component. The unit of temperature must be Kelvin for proper usage. The expected values calculated for the studied system are 120 °C and 135 °C as measured by DSC and DMA, respectively.

The unit of temperature must be Kelvin for proper usage. The expected values calculated for the studied system are 120 °C and 135 °C as measured by DSC and DMA, respectively. Although the PS-PC blends show glass transition temperatures shifted towards the6(1)Fox predicted temperature, no blend studied should “miscible.” Measurements by DMA generally show the printed samples dominated by the PS Tg with, perhaps, a slight peak close to the higher temperature PC Tg. It is interesting that the alternating homopolymer layer samples printed with 150 µm layers shows a tan(δ) peak shifted above temperature and appears simply more dominated by the PC homopolymer layers than either the 300 µm alternating layer sample or either sample printed from twin screw blended filament.beconsideredthe Fox-predictedMeasurements by DSC in Figure 3(b) show two distinct glass transition temperatures.

The extent of shift in the low and high temperature behaviors measured becomes greater with decreasing layer height. This is the hoped-for trend; printing with thinner layers affects a more homogenized blend, or at least spatially distributed morphology. Additionally, it is observed that the temperatures of the twin- screw prepared filament are shifted closer to the Fox-predicted temperature than those from the alternating homopolymer samples at equivalent layer height.glasstransitionImpact on Small Amplitude Oscillatory PropertiesSmall amplitude oscillatory mechanical testing to probe material properties without the sensitivity to surface defects that effects failure properties.

Figure 4 depicts storage modulus as a function of temperature for all samples tested with the 300 µm layer height samples shown in sub-figure a and the 150 µm layer height samples in sub- figure b.is often usedFigure 4: Comparison of storage modulus from printed homopolymers and blends with layer heights of (a) 300 µm and (b) 150 µm. DMA measured in shear mode at 1 Hz; 30 N force amplitude or 2 µm displacement amplitude.Distinguishing behavior between PS (blue) and PC (red) homopolymers are easily observed at both layer heights. The onset of modulus decay, which can be thought of as related to heat deflection temperature, is close to 110 °C for PS but close to 160 °C for PC. Samples printed from 50/50 blended filament (purple) show onset of modulus decay alongside PS,7but the rate of decay is not as severe as the PS homopolymer. The 50/50 blended sample displays a portion of higher temperature stability and stiffness arising from the PC domains.

The 50/50 blended sample displays a portion of higher temperature stability and stiffness arising from the PC domains. Similar behavior is observed for the alternating layers sample at 300 µm layer height (Figure 4a; black); however, the alternating layer sample fabricated at 150 µm layer height (Figure 4b; black) shows extended temperature stability more closely resembling PC. Additionally, the samples with thinner layers show a higher modulus indicative of better interlayer adhesion. This is similar to what Frunzaverde, et al. reported in literature for ultimate tensile properties as a function of layer height.16The authors hypothesize the changes in modulus are a direct result of increased polymer chain mobility at the interface. Abbott, et al. has shown the effect of layer height on temperature profile.17 They report increased time above Tg for thinner layers given similar print conditions.

Abbott, et al. has shown the effect of layer height on temperature profile.17 They report increased time above Tg for thinner layers given similar print conditions. Increased time at temperature (through thinner layers) has a more profound effect on the alternating layer samples, than on those pre-blended at the filament making stage. The interfacial tension at the dissimilar interface can act as an additional driving force for polymer re-entanglement beyond what is typically experienced during homopolymer printing.Impact on Creep RecoveryFigure 5 compares the difference in creep recovery between the samples printed with alternating homopolymer layers and the samples made from 50/50 blended filament at 150 µm and 300 µm layer heights. The data in sub-figure (a) were collected at 128 °C, which is between the tan(δ) peaks measured by DMA and shown in Figure 3.

The data in sub-figure (a) were collected at 128 °C, which is between the tan(δ) peaks measured by DMA and shown in Figure 3. After deformation at elevated temperature, the specimens printed at larger layer height recovered initial dimensions faster and more completely than the 150 µm layer height samples. The samples printed at 300 µm with twin screw blended filament were observed to have the fastest and most complete creep recovery, yet the samples printed at 150 µm with twin screw blended filament were observed to have the slowest and least complete creep recovery. This is consistent for all three load/unload cycles. Since this test was conducted above the Tg of PS and below the Tg of PC, one can expect more permanent deformation of the PS regions due to being highly mobile above Tg.

Since this test was conducted above the Tg of PS and below the Tg of PC, one can expect more permanent deformation of the PS regions due to being highly mobile above Tg. The concentration of PS- and PC-rich regions in the samples with alternating layers likely affords the more consistent behavior regardless of layer height; however, the meso-scale blend morphology introduced at layer interfaces clearly affects the creep recovery of the 50/50 blended samples. The reduction in creep recovery within the time of observation with decreasing layer height (i.e., increased number of interfaces per unit length) implies that the PS plays a major role in the interfacial behavior of the blended samples.8Figure 5: Comparison of creep recovery at (a) 128 °C and (b) 30 °C for samples printed from 50/50 blended filament (purple) and alternating homopolymer layers (black).

Solid lines represent samples printed with 300 µm layer height and dashed lines represent samples printed with 150 µm layer height.Figure 5(b) reports creep recovery at room temperature following the creep recovery experiment at elevated temperature. Here, no meaningful difference is observed between the samples with 300 µm layer height; both exhibit full recovery within the testing window. However, full recovery is not observed for either sample printed at 150 µm layer height, although the sample with alternating homopolymer layers exhibits a greater extent of recovery. The flatline response after returning to room temperature for the 50/50 blended samples printed at 150 µm likely indicates permanent deformation akin to compression set in foams.

The flatline response after returning to room temperature for the 50/50 blended samples printed at 150 µm likely indicates permanent deformation akin to compression set in foams. A comparison against PS and PC printed homopolymer samples will be included as future work.Summary and Future WorkThe presented work explores the unique opportunity afforded by fused filament fabrication (FFF) additive manufacturing (AM) to spatially template a polymer blend in a periodic manner. These samples comprising “alternating homopolymer layers” were compared against the traditional method for preparing polymer blend (i.e., melt-mixing in a twin screw extruder), and the pure polystyrene (PS) and polycarbonate (PC) homopolymer controls. The thermomechanical response of FFF printed samples was measured by dynamic mechanical analysis (DMA) in both oscillatory shear and creep modes in addition to differential scanning calorimetry (DSC).

Glass transition behavior evaluation revealed distinct departure from the expected behavior for a miscible blend based on the Fox equation, thus indicating immiscibility. Trends in DMA behavior interestingly revealed that the sample of alternating 150 µm layers behaved more similar to the PC homopolymer instead of more similar to the traditionally melt-mixed blend. This deviation opens new possibilities for using AM to tailor material performance properties.The difference in response between the samples under creep conditions compared with the trends observed for oscillatory shear conditions demonstrates the importance of evaluating the behavior of AM parts under a variety of load conditions. There have been other discussions in the9literature and broader community concerning the applicability of the ASTM dogbone shape for evaluating AM parts. It is likewise important to investigate how layered objects respond to varied load conditions.

It is likewise important to investigate how layered objects respond to varied load conditions. Polymer chain alignment plays a large role in interfacial behavior and is difficult to erase.Future work should include a more detailed morphology to look at the dispersion of PC and PS in the blended filament. This may be done through either electron microscopy or Raman microscopy. Creep data should be collected for printed versions of PC and PS homopolymers.

This may be done through either electron microscopy or Raman microscopy. Creep data should be collected for printed versions of PC and PS homopolymers. The presented work was able to compare the difference between two methods of preparation for nominally equivalent material compositions, however, a better interpretation can be made in the context of the homopolymer behavior.The differences observed in the presented results arising from AM-tailored distribution of polymers inside a (macro-scale) blended morphology raise questions regarding how layerwise deposition can be exploited towards functionally graded materials or else realizing performance properties not possible to achieve through conventional methods of manufacturing. Since AM is fundamentally a technology reliant on sequentially mating and integrating surfaces, techniques prevalent in adhesion and surface sciences should be considered for increasing understanding of structure-process-morphology-performance relationships.

Conversely, AM can be used as a tool to increase our understanding of surface science since whatever occurs at the layer-to-layer surface is magnified by the number of interfaces and periodicity of interfaces (i.e., layer height) in a part. Such an interdisciplinary pull and push between these two fields will vastly expand our knowledge and increase utility of AM as a viable means for end-use, performance-critical parts.AcknowledgementsThe authors thank Andrew Rhodes who supported sample fabrication and data collection. Blend preparation in the twin-screw extruder was performed by Justin Anderson through Mike Bortner’s Polymer and Composite Manufacturing Laboratory at Virginia Tech.This work was produced by Battelle Savannah River Alliance, LLC under contract number 89303321CEM000080 with the U.S. Department of Energy. Publisher acknowledges the U.S. Government the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).

Gradl3 ,36849, USA2 Department of Mechanical Engineering, Auburn University, Auburn, AL36849, USA 3 NASA Marshall Space Flight Center, Propulsion Department, Huntsville, AL 35812, USACorresponding author: shamsaei@auburn.eduPhone: (334) 844-4839AbstractThe microstructure and mechanical properties of additively manufactured (AM) alloys can be significantly affected by variations in cooling rates, resulting from different process conditions across different additive manufacturing (AM) platforms. Therefore, it is crucial to understand the effect of manufacturing process on the microstructure and mechanical properties of AM Inconel 718. This study examines three AM processes: laser powder bed fusion, laser powder directed energy deposition, and wire arc additive manufacturing.

This study examines three AM processes: laser powder bed fusion, laser powder directed energy deposition, and wire arc additive manufacturing. Results show that fully heat treated laser powder bed fused (L-PBF) and wire arc additively manufactured (WAAM) Inconel 718 specimens exhibit higher strength compared to laser powder directed energy deposited (LP-DED) ones due to finer grain structure in L-PBF and retained dendritic microstructure in WAAM.

The ductility in LP-DED Inconel 718 was slightly higher compared to WAAM and L-PBF due to relatively small carbide size, which causes stress concentration in a small material volume, leading to delayed fracture.Keywords: Additive manufacturing (AM); Laser powder bed fusion (L-PBF); Laser powder directed energy deposition (LP-DED); Wire arc additive manufacturing process (WAAM); Inconel 718IntroductionAdditive manufacturing (AM) offers several advantages over traditional manufacturing such as design flexibility, the ability to fabricate complex geometries, and the capability to manufacture hard-to-machine materials such as Inconel 718 [1]. Different AM technologies have been developed, with laser powder bed fusion (L-PBF), laser powder directed energy deposition (LP-DED), and wire arc additive manufacturing (WAAM) being the most commonly used [2]. Each of these technologies has its unique benefits and limitations.

Each of these technologies has its unique benefits and limitations. For instance, LP-DED and WAAM processes have higher throughput than L-PBF and suitable for larger part sizes, but they have lower accuracy and may not be ideal for intricate features [3].Despite numerous advantages, qualifying and certifying AM parts remain challenging due to the inherent variability in mechanical performance caused by unique micro-/defect- structures. This issue is compounded by differences in cooling rates among technologies, resulting from variations in process parameters, feedstock size, and mass deposition or layer thickness [4,5]. Varied cooling rates have a significant impact on microstructure and mechanical properties [6]. Moreover, the microstructural response to standard heat treatment (HT) can vary significantly across different AM processes [7].

Moreover, the microstructural response to standard heat treatment (HT) can vary significantly across different AM processes [7]. While previous studies have mainly investigated the effects of individual AM processes on microstructure and mechanical properties [8,9], very few have directly compared multiple technologies in terms of microstructural response to standard HT and resulting mechanical properties [3,10]. This study aims to provide a direct comparison of the microstructure and mechanical properties of Inconel 718 manufactured via three commonly used technologies: L-PBF, LP-DED, and WAAM.274Experimental ProcedureThe cylindrical bars of Inconel 718 were fabricated using L-PBF, LP-DED, and WAAM technologies. The chemical composition of powder used in this study for fabrication of L-PBF and LP-DED specimens is exhibited in Table 1. The process parameters employed in L-PBF and LP-DED machines are shown in Table 2.

The process parameters employed in L-PBF and LP-DED machines are shown in Table 2. Note that the process parameters employed in WAAM machine were proprietary and not shown in this article. After fabrication, the specimens in all the manufacturing conditions underwent multi-step HT. This involved stress relief (SR) at 1066°C for 1.5 hr, followed by the hot isostatic pressing (HIP) at 1162°C at 103 MPa for 3 hr, then solution annealing (SA) at 1065°C for 1 hr. Lastly, double aging (DA) which consisted of first step at 760°C for 10 hrs and second step at 650°C for 10 hrs. A schematic illustration of the followed HT is also shown in Figure 1. Following HT, the specimens were machined to the tensile geometry, which had dimensions according to ASTM E8 standard [11] (see Figure 2).For microstructure analysis, small coupons were cut from the tensile specimens and then mounted in cold epoxy resin.

These samples were prepared for scanning electron microscopy (SEM) using a semi-automatic mechanical polisher. The samples were initially ground using sandpapers, then polished on ChemoMet, and finally underwent vibratory finishing to remove any residual layers resulting from grinding or polishing.

All dimensions are in mm.Results and DiscussionThe inverse pole figure (IPF) maps obtained from both non-heat-treated (NHT) and heat treated (HT) conditions of laser powder bed fused (L-PBF), wire arc additively manufactured (WAAM), and laser powder directed energy deposited (LP-DED) samples are presented in Figure 3. In NHT condition, L-PBF Inconel 718 samples exhibited finer grain structure (see Figure 3(a)), which is attributed to their high cooling rates [3]. The grain size distribution revealed that all grains in L-PBF samples were smaller than 100 µm (Figure 3(g)), whereas LP-DED and WAAM samples had predominantly larger grains exceeding 200 µm. After full HT, L-PBF and LP- DED conditions exhibited recrystallization and subsequent grain growth, while WAAM samples retained the as- fabricated characteristics.

After full HT, L-PBF and LP- DED conditions exhibited recrystallization and subsequent grain growth, while WAAM samples retained the as- fabricated characteristics. Moreover, the annealing twins were also observed in LP-DED and L-PBF samples, but not in WAAM ones, further validating that recrystallization did not occur in WAAM samples. The grain size distribution of the HT samples showed that the majority of grains in L-PBF specimens were below 200 µm, while LP-DED and WAAM samples predominantly had grains larger than 300 µm (see Figure 3(h)).276Figure 3 The IPF maps obtained from L-PBF, LP-DED, and WAAM samples in different HT conditions. (g) and (h) exhibit the grain size distribution in NHT and HT conditions, respectively. The black arrows indicate annealing twins.The BSE images obtained from the radial plane (i.e., the plane perpendicular to build direction) of NHT and fully HT specimens in all the processing conditions are presented in Figure 4.

The NHT microstructure in all processing conditions exhibited dendritic microstructure, with finer dendrites observed in L-PBF compared to LP-DED and WAAM samples. Upon full HT, the dendritic characteristics were dissolved in both LP-DED and L-PBF specimens, while WAAM specimens retained the dendritic microstructure. The bright particles within grains and at the grain boundaries are likely Mo-rich carbides, as described in previous studies [12,13]. The size of M6C carbides appeared to be larger in L-PBF and WAAM samples than in LP-DED samples (see insets of Figures 4(b), (d), and (f)). Moreover, the volume fraction of carbides in WAAM appeared to be higher than LP- DED and L-PBF samples.

Moreover, the volume fraction of carbides in WAAM appeared to be higher than LP- DED and L-PBF samples. 277Figure 4 The BSE images obtained from the radial plane of NHT and HT samples in different processing conditions.The tensile properties including yield strength (YS), ultimate tensile strength (UTS), and elongation to failure (EL) are presented in Figure 5 for L-PBF, LP-DED, and WAAM Inconel 718 specimens. Moreover, these properties are compared with those of electron powder bed fused (E-PBF) [9] and wrought counterparts [14]. Interestingly, the wrought parts showed similar strengths and EL to L-PBF and WAAM specimens. Moreover, the YS and UTS of L-PBF and WAAM specimens were higher compared to LP-DED and E-PBF, while EL was slightly lower compared to LP-DED and E-PBF. The difference in strength can be attributed to the finer grain structure in L-PBF and retained as-fabricated characteristics in WAAM.

The difference in strength can be attributed to the finer grain structure in L-PBF and retained as-fabricated characteristics in WAAM. Finer grain size and dendritic microstructures limit dislocation pileup, leading to higher strengths according to Hall-Petch strengthening [15,16]. The higher EL of LP-DED specimens compared to WAAM and L-PBF can be attributed to the smaller size of carbides in LP-DED, which impose stress concentration on smaller material volume. Therefore it is more difficult for smaller carbide to debond from matrix [13]. The tensile fracture surfaces of L-PBF and LP-DED Inconel 718 are shown in Figure 6, exhibiting typical cup and cone fractures with a fibrous region in the center and shear lips around the periphery. Higher magnification images revealed finer dimples in L-PBF specimens, while LP-DED specimens exhibit larger dimples, indicating delayed fracture.

Higher magnification images revealed finer dimples in L-PBF specimens, while LP-DED specimens exhibit larger dimples, indicating delayed fracture. The carbide particles were also observed inside the dimples, validating that carbide debonding governed the fracture behavior in all specimens, regardless of the manufacturing process.278Figure 5 Tensile properties of fully HT and machined L-PBF, LP-DED, WAAM, E-PBF and wrought Inconel 718 specimens.Figure 6 The fracture surfaces of (a)-(b) L-PBF and (c)-(d) LP-DED Inconel 718 specimens. Red arrows indicate fractured particles.279ConclusionsThis study investigated the effect of additive manufacturing process on the microstructure and tensile properties of Inconel 718. The specimens were fabricated using laser powder bed fusion, laser powder directed energy deposition, and wire arc additive manufacturing. All specimens underwent an identical heat treatment schedule.

All specimens underwent an identical heat treatment schedule. The tensile properties generated from this study are compared with those of the wrought and electron powder bed fused counterparts. The following conclusions are drawn from this study:1) The non-heat treated microstructure exhibited dendritic microstructure regardless of the manufacturing condition. Moreover, the size of the grains in laser powder bed fused (L-PBF) Inconel 718 samples was finer compared to laser powder directed energy deposited (LP-DED) and wire arc additively manufactured (WAAM) Inconel 718 ones.2) Upon full heat treatment, dendritic characteristics were not observed in L-PBF and LP-DED samples, while it was observed in WAAM samples.

Moreover, the fully heat treated samples in L-PBF and LP- DED conditions exhibited recrystallization and significant grain growth.3) The grain size of L-PBF Inconel 718 in the fully heat treated condition was finer compared to LP-DED and WAAM ones.4) The fully heat treated and machined specimen in L-PBF and WAAM conditions exhibited higher strength compared to those of LP-DED and electron powder bed fused ones due to finer grain structure in the former and dendritic microstructure in the latter.5) The elongation to failure of LP-DED specimens was slightly higher than WAAM and L-PBF specimens due to smaller carbide particles causing stress concentration in a smaller volume, resulting in delayed microvoid coalescence and fracture.AcknowledgmentThis research is partially supported by the National Aeronautics and Space Administration (NASA) under Cooperative Agreement No. 80MSFC19C0010. This paper describes objective technical results and analysis.

Wixom, MI 48393, USA 5 SLM Solutions Group AG, Lübeck, GermanyKeywords: AlSi7Mg alloy, Laser powder bed fusion, EOS M290 and SLM 280HL systems, Heat treatments, Microindentation hardness (HV), Mechanical properties analysisABSTRACTThe versatility and adaptability of Aluminum F357 (AlSi7Mg) make it a popular material in the aerospace and defense industries. In this study, two different laser powder bed fusion systems, EOS M290, and SLM 280HL were used to create specimens of Aluminum F357. These specimens were subjected to five different heat treatments: As-built, stress relief (SR), hot isostatic pressing (HIP), T6, and HIP+T6) as per ASTM F3318-18 standard. The printed specimens were then reduced to tensile bars through machining and tested for mechanical properties as per ASTM E28 using an MTS Landmark tensile testing system.

The printed specimens were then reduced to tensile bars through machining and tested for mechanical properties as per ASTM E28 using an MTS Landmark tensile testing system. In addition to the mechanical behavior analysis, the study used a JEOL JSM-IT500 SEM to observe and document the fracture produced by the tensile test and a Qness 30 CHD Master+ microhardness testing system to obtain hardness (HV) values of the alloy. The results showed that specimens fabricated in the Z direction had a tendency for higher yield strengths of approximately 225 MPa and although these results were similar between LPBF systems some variances can still be seen. However, these differences between the LPBF systems were observed to be partially mitigated by heat treatments. In conclusion, this study highlights the significance of heat treatment on the mechanical properties of Aluminum F357.

In conclusion, this study highlights the significance of heat treatment on the mechanical properties of Aluminum F357. The results provide valuable information for the aerospace and defense industries to optimize their processes and produce high-quality components. The compatibility of LPBF system fabrication and the mitigation of differences observed between LPBF machines by heat treatments, further demonstrate the potential of this method for producing high-quality Aluminum F357 components.1INTRODUCTIONGenerally, aluminum alloys possess a wide array of desirable engineering characteristics, including low density, high strength, a high strength-to-weight ratio, resistance to corrosion, weldability, and recyclability. Among these properties, strength and related mechanical attributes282such as hardness and elongation can be adjusted to suit various engineering applications through processes like forging, casting, and heat treatment.

These processes result in different microstructures, such as grain sizes, dendrite structures, and interparticle spacings, which directly impact the mechanical properties [1-10]. Additive manufacturing, particularly laser powder bed fusion (LPBF), has introduced additional microstructural features in aluminum alloys due to rapid cooling rates [11]. As a result, new possibilities have emerged for manipulating and selecting mechanical properties through post- process heat treatment, particularly for popular alloys like hypoeutectic AlSi10Mg and AlSi7Mg [12-23]. Heat treatment approaches for LPBF-fabricated AlSiMg alloys include stress relief annealing, T6 treatment (solution heat treatment followed by quenching, cooling, and aging), high-pressure T6 treatment, direct aging, hot isostatic processing (HIP), and various combinations and sequences of these treatments at different temperatures and durations [11,14,17,18,21].

These characteristics of LPBF fabrication and post-process heat treatment have expanded the range of applications for AlSiMg alloys, especially in complex component geometries. Examples include structures for impact and energy absorption, heat exchangers with intricate internal cooling channels, and an increasing number of components in electric vehicles and aerospace systems [12,16,22,23]. This study focuses on the post-process heat treatment of LPBF-fabricated AlSi7Mg (F357) alloy and serves as a follow-up to a comprehensive comparison of heat treatments, residual microstructures, and mechanical properties for AlSi10Mg (A356) alloy, which was recently published [24]. Similar to AlSi10Mg, AlSi7Mg alloy finds wide-ranging applications, as briefly mentioned earlier.

Similar to AlSi10Mg, AlSi7Mg alloy finds wide-ranging applications, as briefly mentioned earlier. The novelty and significance of this research lie in the extensive development of mechanical property data, including 1660 measurements of tensile properties (yield strength, ultimate tensile strength, and elongation) and micro indentation hardness (HV) for components fabricated in the build direction (Z) and the load direction perpendicular to the build direction (XY) in their as-built, aged, heat-treated, and aged states. Additionally, this study compares these mechanical properties for AlSi7Mg components fabricated using two different LPBF systems, which is crucial for standardizing the product.2MATERIALS AND METHODS2.1 Powder feedstocksThe powder utilized in the SLM system was atomized F357 (AlSi7Mg) shown in Table 1; supplied by IMR Metal Technologies (Frankfurt am Main, Germany).

The powder had a particle size distribution of D10: 26.4 μm, D50: 39.7 μm, and D90: 59.7 μm.The powder utilized in the EOS system was atomized F357 (AlSi7Mg) shown in Table 1 and Figure. 1 (a) and (b); provided by Valimet, AM 357C (Stockton, California). A Retsch Camsizer X2 (Haan, Germany) was used in order to study particle size and shape of the powder through dynamic image analysis with a two-camera-system, the Camsizer yields consistent particle size distribution and shape analysis. The stock powder had a particle size distribution of D10: 24.4 μm, D50: 39.3 μm, and D90: 60.4 μm.283Table 1.

The stock powder had a particle size distribution of D10: 24.4 μm, D50: 39.3 μm, and D90: 60.4 μm.283Table 1. Powder chemical composition (wt.-%)AlloyAlSiFeCuMnMgZnTiBeOther elements, eachF357 (EOS) F357 (SLM)BalanceBalance6.5 - 7.57.10.100.080.200.0010.100.0010.40 - 0.700.560.100.0020.04 - 0.200.070.0020.05For further analysis, a JEOL JSM-IT500 Scanning Electron Microscope (SEM) was used to produce images and for analysis which revealed minimal porosity and spherical morphology of the powder, shown in Figure 1.abFigure 1. F357 gas-atomized powder SEM images, a) low magnification, b) high magnification2.2 Laser powder bed fusion systems and process parameters.In accordance with industry best practices, SLM Solutions, based in Lubeck, Germany, employed an SLM 280HL system with a build size measuring 280 x 280 x 365 mm to manufacture specimens.

The F357 alloy was utilized, and the fabrication process involved a laser power of 370 W, a scan speed of 1200 mm/s, and a layer thickness of 30 μm, adhering to the specified parameters.Furthermore, at The University of Texas at El Paso, printed specimens were produced using an EOS M290 system from Kralling, Germany. This particular system features a 400 W Ytterbium fiber laser and a build volume measuring 250 x 250 x 325. After careful consideration of the experimental design, the F357 alloy was employed in the fabrication process. The parameters utilized were a laser power of 370 W, a scan speed of 1300 mm/s, a hatch of 0.13 mm, and a layer thickness of 30 μm.2.3 Thermal processing/heat treatmentAn ASTM standard for heat-treating additively manufactured aluminum F357 does not currently exist. However, in this study on F357, the ASTM F3318 standard was followed.

However, in this study on F357, the ASTM F3318 standard was followed. Previous studies have demonstrated that post-process heat treatments can modify the microstructure and mechanical properties of laser powder bed fusion aluminum F357 parts at room temperature (25 °C). Therefore, this work applied a variety of post-process heat treatments to the F357 specimens. The heat treatments conducted on the F357 specimens included stress relief (SR)284Other elements, total0.15anneals, T6 treatment, hot isostatic pressing (HIP), and combinations of these treatments. Additionally, artificial thermal aging was performed on some specimens to observe changes in microstructure and mechanical properties. Two different aging temperatures (140.5 °C and 177 °C) and three aging times (0 hours, 100 hours, and 1000 hours) were selected.

Two different aging temperatures (140.5 °C and 177 °C) and three aging times (0 hours, 100 hours, and 1000 hours) were selected. The resulting microstructures were characterized using optical metallography, and the mechanical properties were evaluated through room temperature tensile tests and Vickers micro-indentation hardness measurements. With the application of five heat treatments and five different aging conditions, this study presents results for up to 100 different variants of F357. Heat treatments consist in use of temperature and time at temperature to modify the properties of a material, in this case printed aluminum F357. The schedules of the heat treatments consist in the aging of as-built F357 samples, one stress relief treatment (SR1), HIP, T6, and HIP+T6 treatments of the printed samples.

The schedules of the heat treatments consist in the aging of as-built F357 samples, one stress relief treatment (SR1), HIP, T6, and HIP+T6 treatments of the printed samples. ASTM F3318 was used for heat treatment standards shown below:Heat treatments:SR1: 285°C (±14°C) for 120 min (±15 min), air cooled HIP: 100MPa, 515°C (±14°C) for 180 min, inert atmosphere cooled T6: 530°C (±6°C) for 360 min, quenched in water, aged 160°C (±6°C) for 360 min HIP + T6: HIP, followed by T6 processingAging:Times: 0hr, 100hr, 1000hr. Temperatures: 140°C and 177°C.A total of 100 experimental variants were created with these two orientations, times and temperatures as listed in Table 2 and Table 3 below for each of the LPBF systems utilized in this study.

Aging was monitored with external type K thermocouples and recorded by two independent automatic data loggers every 30min to ensure the temperature continue constant at 140°C and 177°C for the 100 and 1000 hours.2.4 Tensile testingAll samples underwent tensile testing using an MTS Landmark servo-hydraulic system, located in Eden Prairie, US, with a force capacity of 10 kN. The system features threaded grips where the specimens are securely held. Additionally, a 30mm axial clip extensometer from MTS was utilized, as depicted in Figure 2. Machining of the samples followed the guidelines outlined in the ASTM E8 standard. To ensure consistent testing conditions, the crosshead displacement speed was set at 0.476 mm/min. For each condition and orientation, the mechanical properties were determined based on the average results obtained from either four or six specimens, as illustrated in Figure 3.

For each condition and orientation, the mechanical properties were determined based on the average results obtained from either four or six specimens, as illustrated in Figure 3. The stress-strain diagram of each specimen was used to calculate its individual mechanical properties. Yield stress measurements were determined using a 0.2% strain offset method. The As-built and HIP conditions comprised four specimens per sample, while the T6, HIP+T6, and SR1 conditions included six specimens per sample.285abFigure 2. a) Tensile test performing sample, b) MTS Landmark servo-hydraulic system2.5 Density measurementsDensity measurements were conducted using the helium gas displacement method employing an Accupyc II 1340 Pycnometer from Norcross, United States. Each experimental variant underwent 10 measurements using the pycnometer. Mass measurements, on the other hand, were obtained utilizing a Sartorius CP124S weight balance manufactured by Sartorius AG in Germany.

Mass measurements, on the other hand, were obtained utilizing a Sartorius CP124S weight balance manufactured by Sartorius AG in Germany. By combining the obtained mass and volume measurements, the density values were calculated.2.6 Hardness measurement[AEG1]Hardness testing was conducted in accordance with ASTM E384-17 standards using a Qatm - Qness 30 CHD Master+ instrument. The Vickers (HV) scale was utilized for the measurements. Samples that were printed in the vertical (Z) direction were evaluated, and measurements were taken from the X, Y, and Z planes. For each sample, five indentations were made on the surface using a load of 100gf. The indentations were spaced at least three millimeters apart.3RESULTS AND DISCUSSION3.1 Microstructure and mechanical property comparisonsFigure 4 compares the light (optical) microscope images for unaged, Z-direction, as-built, SR1 treated, HIP, T6, and T6 + HIP treatment for EOS LPBF fabricated AlSi7Mg alloy components.

The corresponding, representative stress-strain diagrams for each condition (Figure 4 (a) to (e)) are also shown in Figure 4 (f). The initial, as-built microstructure shown in Figure 4(a) is characterized primarily by ~1-micron micro dendritic cells which are altered at the interlayer melt bands. This microstructure is little altered after SR1 anneal at 285 oC (Figure 4(b)). The T6 (at 530 o C); roughly 0.8 melting point (~ 615 o C) treated components as shown in Figure 4(c) completely recrystallizes the dendritic microstructures in Figure 4(a) and (b), creating an average grain size of ~15 microns. This recrystallized grain structure contains a fairly homogeneous distribution of mostly coarse, eutectic Si particles having sizes as large as 5 microns (Figure2864(c)). Following HIP treatment at 515 oC as shown in Figure 4(d), these Si particles have a wider size distribution, and higher particle density. The particles are more globular and the interparticle spacing is ~1 micron.

The particles are more globular and the interparticle spacing is ~1 micron. The corresponding stress-strain diagram shown in Figure 4(f) illustrates a decrease in the YS and UTS, and a corresponding increase in the ductility (elongation) for the HIP treatment in contrast to the as-built, SR, and T6 treated components. These results were also observed for previous studies for LPBF fabrication and heat treatment of AlSi10Mg alloy [24- 26]. The as-built alloy components following HIP + T6 treatment as shown in Figure 4(e) also exhibit a coarse Si particle size smaller than the HIP Si particles in Figure 4(c) as a result of the T6 treatment following the HIP treatment. The particle density and interparticle spacing is similar to those in Figure 4(c) for the T6 treatment.It is useful to examine Figures 5 and 6 simultaneously since they represent the as built and heat treated AlSi7Mg components in Figure 4 following aging for 100 h and 1000 h at 140 oC, respectively.

It is apparent that not only are the component microstructures in Figures 5 and 6 essentially the same, but they are also little changed from those observed in Figure 4; with the exception of the HIP + 1000 h aging at 140 oC (Figure 6 (d)), where the distribution of Si particles is denser. Indeed, the stress-strain diagrams in Figures 5(f) and 6(f) also exhibit essentially the same extremes between the as-built, SR, and HIP treated components as shown in Figure 4 (f). It is notable that the alloy components post-processed by HIP and aged for both 100 h and 1000 h at 177 oC show very dense distributions of the eutectic Si particles: Figures 7(d) and 8(d), respectively. The corresponding stress-strain diagrams shown in Figures 7(f) and 8(f) are also similar and are templates of those stress-strain diagrams shown in Figures 4(f) to 6(f), especially for the HIP treatment.

It might be noted that the eutectic Si particles in all cases are smaller, denser, and have a smaller interparticle spacing for HIP treatment in contrast to either T6 or HIP + T6 treatments: compare Figures 4(d) to 8(d) with Figs 4 to 8(c) and (e). This may represent an anomaly in part since smaller-spaced particles in general raise the strength and lower the ductility (elongation); opposite to the HIP-related stress-strain diagrams especially in contrast to those for the T6 and HIP + T6 treatment diagrams. However, in this case, the very large and irregular Si particles noted for the T6 and HIP + T6 treated components tend to act in a very different way, possibly emulating a duplex structure in connection with the small grains (~ 12-15 microns in contrast to 5-micron Si particles). This would give rise to a higher strength and lower ductility for the T6 and HIP + T6 components shown in the respective stress-strain diagrams [27].

This would give rise to a higher strength and lower ductility for the T6 and HIP + T6 components shown in the respective stress-strain diagrams [27]. It is notable that this phenomenon was also observed in our previous study of LPBF -fabricated and heat treated AlSi10Mg [24], where other related complexities of precipitation hardening were discussed.ab287cdc dc defFigure 4. Microstructure images and corresponding stress-strain diagram for EOS F357 z-axis built and post process heat treatments with no aging. A) As built, b) SR1 c) T6, d) HIP, e) HIP+T6, f) Stress-strain diagram.abcdc dc d288efFigure 5. Microstructure images and corresponding stress-strain diagram for EOS F357 z-axis built and post process heat treatments aged for 100h at 140°C. a) As built, b) SR, c) T6, d) HIP, e) HIP+T6, f) Stress-strain diagram.abcdc dc d289efFigure 6.

a) As built, b) SR, c) T6, d) HIP, e) HIP+T6, f) Stress-strain diagram.abcdc dc d289efFigure 6. Microstructure images and corresponding stress-strain diagram for EOS F357 z-axis built and post process heat treatments aged for 1000h at 140°C. a) As built, b) SR1 c) T6, d) HIP, e) HIP+T6, f) Stress-strain diagram.While Figures 4 to 8 only represent the range of microstructures and tensile (stress-strain) properties observed for EOS system-fabricated and post-process heat treated AlSi7Mg alloy components with and without aging treatments corresponding to the Z or build direction, components similarly fabricated and post-processed in the XY direction (perpendicular to the build direction as shown in Figure 3) produced essentially identical sequences of microstructures and corresponding stress-strain diagrams.

In addition, components fabricated in the SLM LPBF system also produced the same microstructures and stress-strain diagram trends as those observed for the EOS system fabrications in both the Z and XY directions (Figure 3). Figure 9 provides support for this contention by comparing the stress-strain diagrams for both the EOS and SLM system components fabricated in both loading directions (Z and XY) and heat treated without aging. It is notable in Figure 9 that the extremes in the stress-strain diagrams show uniformly that the as-built components have the highest strength (YS and UTS) and lowest elongation to fracture while the HIP components exhibit the lowest strength and highest elongation to fracture, consistent with previous studies [24-26]. Note also that the trends of the stress-strain diagrams in Figure 9 are essentially the same for both the EOS system and SLM system fabricated AlSi7Mg components.ab290cdc dc defFigure 7.

Microstructure images and corresponding stress-strain diagram for EOS F357 z-axis built and post process heat treatments aged for 1000h at 177°C. a) As built, b) SR, c) T6, d) HIP, e) HIP+T6, f) Stress-strain diagram.abcdFigure 9 Characteristic curves for a) EOS printed parts in Z orientation, b) SLM printed parts in Z orientation, c) EOS printed parts in XY orientation and d) SLM printed parts in XY orientation.Figure 10 represents alternative comparisons of stress-strain diagrams for components fabricated in the EOS system and aged along the left-hand portion of the figure (Figure10 (a), (c), (e), and (g), in contrast to corresponding SLM-fabricated components along the right-hand column (Figure 10 (b), (d), (f), and (h)), respectively. Here again there is a similarity in the trends of the individual stress-strain curves.

Here again there is a similarity in the trends of the individual stress-strain curves. A notable difference between these diagrams is their shape as292they tend to fracture for the SR-treated components in Fig 10 (e) and (g) for the EOS system and Figure 10 (f) and (h) for the SLM system; representing tensile testing in the Z and XY directions, respectively. This steep decline in stress over a narrow strain window is indicative of rapid thermal softening or necking as a result of a rapid increase in void fraction leading to failure. This phenomenon has been discussed in some detail by Pineau, et al. [28] and Wcislik and Pala [29] for metal failure in general.Figures 11 and 12 illustrate that there is a preponderance of this steep softening for HIPed and aged tensile components fabricated in the Z and XY directions in both the EOS and SLM systems; especially notable in the XY direction as shown in Figure 12 (a) and (b).

These results indicate that in some applications, components of AlSi7Mg alloy fabricated in the XY direction in either the EOS or SLM systems may pose some concerns after HIP and adding, which create a broader range of lower strengths in contrast to other heat treatments.abcd293efghFigure 10. Characteristic curves for printed parts on a) EOS Z as built, b) SLM Z as built, c) EOS XY as built, d) SLM XY as built, e) EOS Z stress relief, f) SLM Z stress relief, g) EOS XY stress relief, h) SLM XY stress reliefabcd294efFigure 11. Characteristic curves for printed parts on a) EOS Z HIP, b) SLM Z HIP, c) EOS Z T6, d) SLM Z T6, e) EOS Z HIP+T6, f) SLM Z HIP+T6.abcdef295Figure 12.

Tables 4 to 7 list the average, measured mechanical property values (YS, UTS and elongation); including densities and microindentation hardness (HV) values for the various components. It can be observed in Tables 4 to 7, and as noted in the various stress-strain diagram comparisons, that the as-builtTable 4. Mechanical properties and related data for as built, SR, HIP, T6 and HIP+T6 on EOS Z build directionEOS (Z) Heat Treatment Aging Condition Avg. YS (Mpa) Avg.

Additionally, the XY-built component YS and elongations are generally larger than corresponding as-built Z direction components, while the HIP component values are little changed, indicating more homogenized microstructures and corresponding mechanical properties.

This applies to the components fabricated in either the EOS or SLM systems, Confirming that AlSi7Mg components fabricated, and heat treated from either LPBF system do not differ significantly in their mechanical properties.It is also of interest to observe in Tables 4 to 7 that aging at 177 oC for 1000 h decreased both the yield stress and corresponding micro indentation hardness (HV) in contrast to the unaged condition for all as-built, HIPed, T6 and HIP + T6 treated components for both loading directions (Z and XY) for both LPBF systems (EOS and SLM); the exception being the SR- treated components where both the YS and HV values increased from the unaged condition.It can also be observed that the trends in the associated micro indentation hardness (HV) values correspond to the YS values: high YS values for the as-built components correspond to high HV values while the low YS values for HIP components correspond the low HV values for these components.

It is observed in Tables 5 to 7 that products fabricated in the XY direction have YS Values ~ 12% higher than those fabricated in the Z direction. In contrast, elongations for HIP components are generally similar to either Z or XY builds, although the highest elongations of ~ 40 % occur uniformly for HIP + aging at 177 oC for 1000 h.297Heat Treatment As BuiltSRHIPT6HIP + T6Table 5. Mechanical properties and related data for as built, SR, HIP, T6 and HIP+T6 on SLM Z build direction SLM (Z)Aging Condition 0hr 100hr 140°C 1000hr 140°C 100hr 177°C 1000hr 177°C 0hr 100hr 140°C 1000hr 140°C 100hr 177°C 1000hr 177°C 0hr 100hr 140°C 1000hr 140°C 100hr 177°C 1000hr 177°C 0hr 100hr 140°C 1000hr 140°C 100hr 177°C 1000hr 177°C 0hr 100hr 140°C 1000hr 140°C 100hr 177°C 1000hr 177°CAvg. YS (Mpa)263 271 202 182 183 156 156 149 158 177 85 88 93 88 80 267 292 268 201 102 199 314 278 212 111Avg.

YS (Mpa)295 283 215 196 188 155 164 148 163 180 125 90 95 88 82 260 295 258 197 103 228 309 286 210 111Avg. UTS (Mpa)412 392 322 303 285 241 234 226 245 254 181 139 137 129 120 311 334 297 235 156 281 352 326 247 162Avg Elongation (%)13.1 11.8 14.4 14.7 16.5 19.3 18.0 21.4 19.7 17.5 17.6 36.0 32.6 35.7 39.8 12.9 10.4 8.9 14.1 19.8 16.5 15.9 13.9 14.3 23.1Density (g/cm3) 2.65 2.65 2.66 2.65 2.67 2.65 2.65 2.65 2.65 2.65 2.65 2.57 2.66 2.56 2.65 2.63 2.56 2.61 2.59 2.67 2.66 2.66 2.67 2.66 2.65It should also be mentioned that the measured densities shown in Tables 4-7 vary from 2.53 g/cm3 to 2.67 g/cm3, but this spread in measured densities does not appear to have any systematic heat treatment basis and would appear to be primarily measurement error. There is correspondingly no correlation between the measured densities and the associated mechanical properties, including the micro indentation hardness.

There is correspondingly no correlation between the measured densities and the associated mechanical properties, including the micro indentation hardness. It might be noted that solution treatments of AlSi10Mg at 540 oC reduced the density as a result of diffusion of dissolved hydrogen and increasing gas porosity in a study by Girelli, et al. [30]. However, there is no evidence of these phenomena in the present study.While Tables 4 to 7 provide a comprehensive matrix for design strategies and mechanical property selection for LPBF-fabricated AlSi7Mg alloy products, the trending of properties, particularly mechanical properties, are difficult to compare, even including the stress-strain diagram comparisons shown in Figures 4 to 12.

While it may seem overly redundant, we have300Hardness (HV)133 130 101 95 93 79 81 74 80 90 49 49 49 47 43 107 120 108 81 56 65 122 111 89 58prepared a series of comparative and systematic bar graph summaries illustrating the heat- treatment regimens and corresponding aging treatments for AlSi7Mg alloy fabricated in both the EOS and SLM LPBF systems, and in both the Z and XY loading directions, relative to the build direction. These are shown in Figures 13 and 14. It is interesting to observe that these two comparative figures, representing each LPBF system (EOS and SLM) are essentially templates of one another; illustrating essentially the same trending as discussed extensively above. It can also be noted that Figure 13 for the EOS system is a template for the same heat treatment schedule trends determined for LPBF fabrication of AlSi10Mg alloy components reported in Figure 5 of our previous article [24].

It is also observed that the as-built and as-built and aged components in Figures 13 and 14 exhibit a reversed trend for YS and UTS versus elongation, in contrast to HIP and aged components: higher YS and UTS exhibit lower elongations while lower YS and UTS exhibit higher elongations. It will be noted in retrospect that the Introduction in this study illustrated a range of prior research work describing post-process heat treatments for LPBF fabrication of AlSi10Mg alloy components [11,15,17,18-20,24]. However, far fewer studies have been reported for AlSi7Mg (F357) alloy [10,11]. Vanzetti, et al. [31] have recently described short heat treatments for F357 (AlSi7Mg) alloy processed by LPBF, and involving T6 treatment and direct aging. They observed an optimized yield stress of 308 MPa and an elongation of only 3.9 %; nearly 1/3 of the present as-built elongations having similar yield stress values. Similarly, Fiocchi, et al.

They observed an optimized yield stress of 308 MPa and an elongation of only 3.9 %; nearly 1/3 of the present as-built elongations having similar yield stress values. Similarly, Fiocchi, et al. [32] in their recent review of heat treatment of LPBF-fabricated aluminum alloys noted that thermal treatment of AlSi7Mg alloy actually reduced the as-built elongation from 7.6 % to 4.7 % at a relatively constant yield stress of 257 MPa; less than half the elongation at the same yield stress in the present study (see Tables 4 to 7). In a recent paper by Pezda [10] dealing with the optimization of T6 heat treatment parameters of cast AlSi7Mg alloy where elongations increased by 250 % at aging above 300oC; characteristic of HIP treatment and aging as shown in the current study Correspondingly, there are few prior reports involving HIP of LPBF-fabricated AlSi7Mg alloy. A recent study by Oliveira de Menezes, et al.

A recent study by Oliveira de Menezes, et al. [33] for LPBF fabrication of AlSi7Mg alloy examined the effect of heat treatment parameters and loading orientation on the tensile properties. This study pointed out that while there is a notable anisotropy in mechanical properties between components built parallel and perpendicular to the build direction (Z), this can be reduced or eliminated by heat treatment (such as T6 temper) as observed for numerous other aluminum alloys [11-13, 22,24], as well as many other metals and alloys [34,35]. Indeed, it is observed in the current study (Tables 4-7) that AlSi7Mg alloy components built in both the EOS and SLM LPBF systems exhibit yield stress values (YS) in the Z (vertical) and XY (horizontal) loading directions of 225 MPa and 263 MPa, and 257 MPa and 295 MPa, respectively: an anisotropy of ~ 16%.

Comparative bar graphs showing mechanical properties for F357 on SLM printer, under different aging conditions for the different heat treatments and built in Z and XY orientation, a), d) Yield strength, b), e) UTS and c), f) Elongation; notice strength decreases with aging in a), b), d) and e), while elongation increases in c) and f) for T6, HIP and HIP+T6.303HIP+T6HIP+T6HIP+T64SUMMARY AND CONCLUSIONSThis research program is a comprehensive investigation into the fabrication and heat treatment of AlSi7Mg (F357) alloy using laser powder bed fusion (LPBF). Tensile specimens were fabricated in the Z (build) and XY loading directions on two LPBF systems: EOS M290 and SLM 280HL. The components underwent aging for 100 hours and 1000 hours at temperatures of 140°C and 177°C, respectively.

The components underwent aging for 100 hours and 1000 hours at temperatures of 140°C and 177°C, respectively. Heat treatment processes included stress relief (SR1) anneal, hot isostatic pressing (HIP), T6 treatment, and HIP + T6 treatment.A total of 1660 measurements were conducted, including yield stress (YS), ultimate tensile strength (UTS), elongation, and micro indentation hardness (HV). Density measurements were also performed. The mechanical properties exhibited anisotropy between the Z and XY loading directions for each LPBF system. For example, the unaged YS of EOS (Z) was 225 MPa, while EOS (XY) had a YS of 257 MPa. Similarly, the unaged YS of SLM (Z) was 263 MPa, whereas SLM (XY) had a YS of 295 MPa.The highest YS values were consistently observed for the T6 treatment + 100-hour aging at 140°C. For example, the EOS (Z) and (XY) components had YS values of 291 MPa and 292 MPa, respectively.

Similarly, the SLM (Z) and (XY) components had YS values of 292 MPa and 295 MPa, respectively.The micro indentation hardness (HV) measurements were consistent with EOS (Z) and EOS (XY) exhibiting hardness values of HV 120 and HV 122, respectively, while SLM (Z) and SLM (XY) had hardness values of HV 122 and HV 133, respectively.These findings demonstrate the influence of heat treatment and aging on the mechanical properties of LPBF-fabricated AlSi7Mg (F357) alloy components, hence we can conclude the following:Although only two different LPBF systems (EOS and SLM) were compared in this study, it would appear that machine differences do not significantly alter AlSi7Mg alloy products produced, and any inhomogeneities and anisotropies are erased as a result of heat treatment, especially HIP, T6 and HIP + T6 treatments.As-built AlSi7Mg alloy products fabricated in the build (Z) direction varied in yield stress from 225 MPa to 295 MPa in this study; with corresponding elongation variations 0f 7 % to 17 %.

These properties reflect the most economical and process efficient strategy.While HIP treatment can produce a homogenized microstructure even without aging, the resulting yield stress values were observed to vary from 85 MPa to 87 MPa for Z-direction fabricated components; with corresponding elongations varying from 25 % to 31 %. The highest elongation was observed for HIP + 1000 h aging at 177 oC: 42 %.Bar graph representations of the tensile properties (YS, UTS, elongation) for both the LPBF systems heat treated and heat treated + aging (as shown in Figures 13 and 14) were templates of one another.304REFERENCES[1] Moller H, Govender G, Stumpf WE, Pistorius PC. Comparison of heat treatment response of semisolid metal processed alloys A356 and A357. Int J Cast Met Res 2010; 23: 37-43. [2] Totten GE, editor. Heat treating aluminum and its alloys. Materials Park, OH: ASM International; 2016. [3] Totten GE, Tiryakioglu M, Kessler O. editors.

Geometry of dimples and its correlation with mechanical properties ofAustenitic stainless steels. Scripta Mater 2008; 59(9): 1014-1017.[40] Hilders OA, Zambrano N, Caballero R. Microstructure, strength and fracture topography in AISI 304L stainless steel as seen through a fractal approach and Hall-Petch law. Int J Metals2015; 10: 624653. [41] Manjunath GK, Preetham Kumar GV, Bhat KU. Tensile properties and tensile fracture characteristics of cast Al-Zn-Mg alloys processed by equal channel angular pressing. Trans Indian Inst Met 2017; published online Feb. 17, 10 pp.ACKNOWLEDGEMENTS Primary support for this research was provided by grant number NOAID20210005 from AFRL (Air Force Research Laboratory) as well as strategic investments via discretionary UTEP Keck Center funds and the Mr. and Mrs. MacIntosh Murchison Chair I in Engineering Endowment at UTEP. The research described here was performed at The University of Texas at El Paso (UTEP) within the W.M.

and Mrs. MacIntosh Murchison Chair I in Engineering Endowment at UTEP. The research described here was performed at The University of Texas at El Paso (UTEP) within the W.M. Keck Center for 3D Innovation (Keck Center). The authors are grateful to Jorge Merino for its guidance in the writing of this paper. The views and opinions expressed in this article are those of the authors and do not necessarily reflect the official opinion or policy of any agency of the US government (if applicable).307.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceA NOVEL COATING METHOD USED TO ENABLE MULTILAYER STRUCTURES WITH MICROSCALE SELECTIVE LASER SINTERINGA. Liao*, D. Behera*, M. A. Cullinan*Walker Department of Mechanical Engineering, The University of Texas at Austin, Austin, TX 78712AbstractThe microscale selective laser sintering process (µSLS) is an additive manufacturing technique that enables the creation of metal features with sub-5 µm in-plane resolution. In this process, a layer of metal nanoparticle ink is deposited onto a substrate and positioned beneath an optical subsystem with a nanopositioning stage. Using a digital micromirror device, a laser is spatially modulated to selectively heat up particles in desired regions to cause sintering. The substrate is then moved to a coating station where a new layer of nanoparticle ink is applied atop the sintered features.

The substrate is then moved to a coating station where a new layer of nanoparticle ink is applied atop the sintered features. Initially, the slot-die coating process was adopted as the recoating method for this technique. However, due to challenges with depositing consistent ink thickness across the recoated part and limitations with the minimum layer thickness achievable, a new approach inspired by blade coating has been developed to achieve layer thicknesses of less than 1 µm.IntroductionThe aim of developing the microscale selective laser sintering process is to manufacture three-dimensional components with sub-5 µm features, all while achieving a throughput of 60 mm3/hr [1]. To create a single layer of a part the machine must heat a particle bed at a designated location, apply a new layer of nanoparticles onto the sintered part, and accurately position the components between successively added layers.

To carry out these tasks, the current version of the machine comprises three primary subsystems: an optical sintering system, global and nanopositioning systems, and a nanoparticle ink dispensing system as shown below. These processes are iteratively executed to eventually build up a three-dimensional part.Figure 1: Diagram showing the three subsystems of the microscale selective laser sintering machine and their components [1].308BackgroundBecause microscale selective laser sintering (µ-SLS) of metals adheres to similar principles as its larger-scale counterpart, achieving uniform deposition of nanoparticles onto previously sintered layers is crucial. Instead of utilizing micron-scale particles as seen in traditional metal powder bed fusion (PBF), µ-SLS employs metal nanoparticles to enhance process resolution.

Instead of utilizing micron-scale particles as seen in traditional metal powder bed fusion (PBF), µ-SLS employs metal nanoparticles to enhance process resolution. However, obtaining repeatable, densely packed layers with nanoparticles presents challenges due to their elevated surface energy and susceptibility to excessive oxidation [2]. Previous sintered structures created with nanopowders had high porosity, subpar physical attributes, and low electrical conductivity [3]. This prompted investigation into using nanoparticle inks, which consist of copper nanoparticles dispersed in an organic solvent. Initially, the slot die coating process was chosen for this process, however the films deposited by this process were too thick, and this work investigated using a different meniscus guided coating process.Blade coating was of particular interest due to ease of setup and significantly reduced volume of ink in , Landau Levich coating theorythe meniscus.

Under fluid flow condition with low capillary number ( can be used to predict the thickness of films deposited by meniscus guided processes. Capillary number is defined as the ratio between viscous drag forces and surface tension forces and can be expressed as in Equation 1 where is the surface energy of the fluid.𝐶𝐶𝐶𝐶 ≪ 1)is the viscosity of the fluid,is the velocity of the fluid meniscus, and𝜂𝜂𝜈𝜈𝛾𝛾(1)𝜂𝜂𝜈𝜈 𝛾𝛾𝐶𝐶𝐶𝐶 =The solution derived by Le Berre et al.

[2] for a is provided below in Equation 2.(2)𝑡𝑡𝑤𝑤𝑤𝑤𝑤𝑤 = 1.34𝐿𝐿(𝐶𝐶𝐶𝐶)2/3𝐻𝐻(3)2𝐿𝐿 =𝐻𝐻 2𝜅𝜅cos 𝜃𝜃1 + cos 𝜃𝜃2 −−2The geometry of the meniscus in the flow coating process is shown in the figure below, labeling wherethe constants in Equation 3 are from.Figure 2: Diagram of the meniscus geometry in the flowcoating process [3]This solution for wet film thicknesses deposited by the flowcoating process will be compared against theexperimental results by obtaining meniscus geometry parameters.309Blade Coating Apparatus DesignThe coating apparatus should be able to apply extremely thin layers of metal nanoparticle ink (less than1 µm in thickness) onto sintered metal surfaces from the preceding layers. The coating process window, thickness, and uniformity for the previously used slot die coating procedure, are influenced by various process parameters.

The coating process window, thickness, and uniformity for the previously used slot die coating procedure, are influenced by various process parameters. Given that this approach is still influenced by the meniscus, we aim to have similar control over process parameters as those employed in slot die coating. The coating device should have the ability to control the coating gap, coating speed, and volume of deposited ink. The tool that was developed was able to automate the blade positioning, ink dosage, and substrate translation steps required to deposit a thin film with the fluid entrained by the meniscus.Figure 3: a) CAD model of the assembled coating device. b) Physical embodiment of the coating device. c) Zoomed in view of the custom designed 3D printed parts that hold blade, distance sensor and pipette tip.The coating blade height was controlled using a 5-phase stepper motor (Oriental Motor) to drive a screw driven linear stage (Newport).

The minimum step of the motor and the pitch of the screw results in a resolution of 1 μm, allowing for precise positioning of the blade relative to the substrate. This distance is measured using a laser triangulation distance sensor (Keyence), and an open loop stepping algorithm is used to position the blade at various heights.Ink is delivered to the meniscus by using a syringe pump (Kloehn/IMI Norgren) with a 5 mL zero dead volume tip syringe. It is plunged by a brushless DC motor with a resolution of 24000 pulses per revolution, and lead screw with 1mm pitch, allowing us to achieve a volumetric resolution of approximately 300 nL.To provide the substrate translation, the nanopositioning stage was used. The stage is translated along the axis aligned with the coating direction using a voice coil and a 1:1 current amplifier. To keep track of its position, an Attocube FPS 3010 interferometer is employed, featuring a Thorlabs silver mirror target attached to a 3D-printed bracket.

To keep track of its position, an Attocube FPS 3010 interferometer is employed, featuring a Thorlabs silver mirror target attached to a 3D-printed bracket. The displacement of the stage in this axis is controlled by a PID controller with the output current of the voice coil actuator. With custom trajectory generation, we can precisely control the speed of the substrate for the coating operation.These are all automated within LabVIEW and a schematic showing the process and the order of thesteps is shown below.310Figure 4: Sequence of the coating steps automated by the prototype coating tool. a) Begins with blade lifted off substrate surface pipette tip away from gap. b) Servo actuates pipette tip to substrate gap. c) Syringe pump forces ink through tubing that exits at the pipette tip, forming meniscus. d) Servo actuates pipette tip off glass substrate. e) Substrate travels with ink entrained in meniscus.

c) Syringe pump forces ink through tubing that exits at the pipette tip, forming meniscus. d) Servo actuates pipette tip off glass substrate. e) Substrate travels with ink entrained in meniscus. f) Blade lifts away from substrate and breaks meniscus.Characterization of Single Layer CoatingTo control the thickness of the ink deposited, we want to determine relationship between coating speed and wet film thickness. The coating gap (100 μm), and volume of ink (3 μL) in the meniscus were held constant and the coating speed in these experiments ranged between 0.5 – 3 mm/s. To compare against the coating theory, images of the blade, substrate, and meniscus were taken to obtain the geometric constants used in Equation 3.

Figure 5 shows the side view of the meniscus geometry and the relevant lengths and angles required to estimate the wet film thickness using Landau Levich coating theory.Figure 5: Image of the meniscus used to derive geometric parameters used in the calculation of predicted wet film thicknesses.311Given that the capillary number exhibits a direct proportionality to velocity (as per Equation 1), andthickness follows a power-law relationship with the capillary number (as per Equation 2), the velocity should also exhibit a power-law relationship with the thickness of the ink.Figure 6: Plot showing predicted and measured wet film thicknesses achieved by the meniscus geometry of the blade coating device.In our experiments, we observed that although the coating thickness did increase with increasing velocity, the curve fit had a lower power law exponent of 0.391 instead of the expected 0.667 from Equation 2.

Additionally, the magnitude of the thickness is also around 5 times higher than predicted as opposed to the 2 – 3 times seen in the literature for this meniscus geometry.Multilayer Coating PerformanceThe fabrication of multilayer samples using the slot-die coating process presented difficulties, includinginconsistent features, unmanageable part thicknesses, and insufficient adhesion of deposited material to the substrate and to previous layers.

This coating method struggled with the deposition of thin ink layers, and when coupled with the laser's limited sintering depth, it resulted in the creation of easily discernible bands of sintered and unsintered nanoparticles, as illustrated in the figure below.312Figure 7: Cross section of multilayer coating process with the previous slot die coating process.With the new coating process able to deposit submicron layers of ink (~ 0.3 μm), we can create structures that are fully sintered with no banding visible as shown in Figure 8 below. This leads to significantly improved interlayer adhesion which improves the final part strength and properties.Figure 8: Cross section of the multilayer coating process using the newly designed blade coating process.To test the stacking of multiple sintering regions, an 18-layer staircase structure was created using thelaser masks below.

Each of the images was projected by the DMD to heat up the particles in the white region of the mask, with each step using a smaller area to create a pyramid-like shape.313Figure 9: Three masks used to generate staircase pattern, and their respective layer numbers.Figure 10: Staircase feature showing 3D capability of the μ-SLS process. Red line on left image shows where height profile was measured.Each "staircase" step is comprised six consecutive layers of the same mask, which should roughly give acomponent with an approximate height of 5.8 μm. Considering that we anticipate roughly 0.3 μm per recoat in the coating process, the expected component height was projected to be around 5.4 μm.

Considering that we anticipate roughly 0.3 μm per recoat in the coating process, the expected component height was projected to be around 5.4 μm. The envisioned profile was graphed and compared to the real profile along the center of the rectangle in the figure below.Figure 11: Averaged line profile across the center of the stacked pyramid structure.314The step heights closely align with the anticipated increment in layer height brought about by the bladecoating process. In summary, we can see that this technique shows significantly improved vertical resolution compared to what was previously observed in the multilayer performance of the slot die coating process.ConclusionThe objective of this work was to develop and evaluate a novel recoating process for the microscaleselective laser sintering process to enable the creation of multi-layered structures.

After an initial assessment of film deposition thickness, evaluation of improved sintering depth, and creating multilayer parts, it is evident that the blade recoating process has the capacity to deposit submicron layer thicknesses in a manner suitable for manufacturing multi-layered components. This process demonstrated reliable repeatability for components with fewer than 20 layers. However, achieving more precise control over layer thickness necessitates comprehensive characterization and a deeper exploration of the fundamental physics involved. This entails quantifying and comprehending the impact of coating operation parameters beyond just coating speed, including coating gap, ink volume, and blade angles. As the number of layers increases, a thorough understanding of these parameters becomes essential to maintain the recoating process's repeatability and prevent cumulative errors.315References[1] N. K. Roy, D. Behera, O. G. Dibua, C. S. Foong, and M. A.

The laser head is only partly inserted into the chamber through the cover. This enables a small sized chamber and allows a quick filling with argon. Since the chamber has a low leakage, only small amounts of argon (5 l/min) are needed to maintain a sufficient welding atmosphere with less than 300 ppm oxygen. For large sized parts, the chamber can be repositioned on the substrate. It has flexible parts which can be fit to the already welded structures that otherwise would prevent the chamber from being put flat on the substrate. The limited build space inside the chamber requires a new welding strategy, which is suggested.IntroductionAdditive Manufacturing (AM) has a promising economical potential for aerospace applications, especially for bigger titanium parts where sometimes over 90 percent of the material is taken off by milling. In these cases, many of the expensive alloys like Ti-6Al-4V can be saved.

In these cases, many of the expensive alloys like Ti-6Al-4V can be saved. Since titanium has a high affinity to oxygen at elevated temperatures, a shielding gas atmosphere has to be applied during welding until it has cooled below 430°C (800°F) [1,2] to protect it from oxidation. Oxygen is an α-phase stabilizer like aluminum in contrary to the vanadium, which is a β-phase stabilizer. If present at elevated temperatures during welding, it leads to the so-called α-casing, an oxidation that can be seen by the discoloration of the surface. The more oxygen the titanium picks up, the more α- phase will be present, leading to an increase of brittleness and a decrease of ductility.In general, there are three different ways to protect titanium from oxidation. The use of a vacuum chamber is one possibility. This technique is mostly used when welding with an electron beam, since the process requires a vacuum. Establishing a vacuum is a rather slow process, especially in a big chamber.

This technique is mostly used when welding with an electron beam, since the process requires a vacuum. Establishing a vacuum is a rather slow process, especially in a big chamber. For this reason, a protection gas chamber is more common when welding metals with high affinity to oxygen or other gases with impact on the microstructure. A literature research reveals a broad range regarding the level of oxygen content in the argon atmosphere and seems to be dependent on the type of AM process. Ding et al. [3] found that the hardness of WAAM (wire arc additive manufacturing) welded Ti-6Al-4V components were not affected by an oxygen content of 4000 ppm in the argon atmosphere. Therefore, they selected 2000 ppm as a conservative environmental contamination level. Caballero et al. [4] also investigated the effect of oxygen in TI-6Al-4V with WAAM. They found that tensile properties are not compromised by an increase of oxygen content in the shielding gas of up to 4000 ppm.

[4] also investigated the effect of oxygen in TI-6Al-4V with WAAM. They found that tensile properties are not compromised by an increase of oxygen content in the shielding gas of up to 4000 ppm. High temperatures and exposure times seem to have a greater effect on oxidation than the oxygen content in the shielding environment. They also found that different oxygen contents in the wire do not have a significant effect on the thickness of the alpha case. Halisch et al. [5] used a WAAM process in a protection gas chamber to evaluate the influence of oxygen content in the shielding on the mechanical properties of TI-Al6-V6 parts. They concluded that with oxygen levels of up to 6000 ppm the requirements on tensile properties for aerospace parts could be reached, even though yield and tensile strength increases with the oxygen level while the elongation decreases. Gushchina et al.

Gushchina et al. [6] investigated the influence of the oxygen levels in the protection gas atmosphere during additive manufacturing of Ti-6Al-4V using a laser powder DED process. Growth of oxygen concentration until 0.25% O2 led to change of mechanical characteristics. The best mechanical properties by strength to ductility ratio were produced with 0.25% oxygen concentration in chamber. The author did not specify1462whether weight or atomic percentage is meant. Since weight percentage would result in 1000 ppm oxygen content this is presumed by the author of this study. Chekir [7] printed Ti-6Al-4V parts using a laser wire process in a process chamber with below 60 ppm oxygen in the argon environment to prevent excessive oxidation. Dai et al. [8] kept the oxygen content below 200 ppm to impede oxidation for wire laser AM with an oscillating laser beam. Gibson et al. [9] used a laser hot-wire process to print Ti-6Al-4V parts.

Dai et al. [8] kept the oxygen content below 200 ppm to impede oxidation for wire laser AM with an oscillating laser beam. Gibson et al. [9] used a laser hot-wire process to print Ti-6Al-4V parts. A chamber with a flexible bag on top was used to maintain an atmosphere with oxygen levels under 300 ppm during metal deposition. The printed part nevertheless exhibited yellow and blue oxide in some regions. Velasco-Castro et al. [10] report an increase in hardness and a decrease in ductility of titanium parts manufactured by LPBF (laser powder bed fusion) in an inert gas atmosphere with 1000 ppm oxygen content. Emminghaus [11] showed that the most influential factor concerning the mean hardness in Ti-6Al-4V parts manufactured with LPBF is the residual oxygen content. Pauzon et al. [12] found that an oxygen control to 100 ppm is required to impede these effects in LPBF. Tapoglou et al.

Pauzon et al. [12] found that an oxygen control to 100 ppm is required to impede these effects in LPBF. Tapoglou et al. [13] realized a protection gas atmosphere for additive manufacturing of titanium using a tub on the bottom and a plastic bag on the top. Argon flow rates of 50 l/min were applied to the chamber reaching oxygen levels of 100 ppm. The AM process used Ti-6Al-4V powder with a flow rate of up to 12 g/min (0.72 kg/h) at 2000 W laser power. This results in a high use of argon and only 0.24 g titanium per liter argon can be build. Yu et a. [14] used a laser powder DED process to build Ti6Al4V parts. The parts were built without the use of an inert gas chamber but with local protection using an argon gas flow provided by the powder nozzle. The maximum powder flow was 1.98 g/min or 119 g/h at 570 W laser power with a protection-gas flow-rate of 10.5 l/min. This means a relatively low heat and material input compared to laser wire processes.

This means a relatively low heat and material input compared to laser wire processes. The process uses high amounts of argon. Only 0.189 g per liter argon can be build. As mentioned before, AM of titanium has a promising economical potential. This can be increased by reducing the amount of argon used for the building process.Experimental and ResultsTo reduce the argon consumption needed for a sound protection-gas-atmosphere when manufacturing bigger parts, a new chamber concept (Figure 1) has been tested. The inner volume of the chamber has a volume of 5.5 l and encloses only a small part of the welding head. In order to manufacture bigger parts, the chamber can be moved successively along areas to weld. The cover of the chamber is slid to the sides by the movement of the laser head. Since it is bigger than the chamber itself, the chamber will not open during this process. A cylinder, which is attached to the laser head, allows vertical motions.

Since it is bigger than the chamber itself, the chamber will not open during this process. A cylinder, which is attached to the laser head, allows vertical motions. When placing the chamber in a new location, build-up parts may block the chamber from being put flat on the surface, hindering the establishment of an intact protection- gas-atmosphere. Therefore seals are installed, consisting of thin metal rods, which can individually be moved vertical and allow to adapt to obstacles.Figure 1: Rendering of the protection-gas chamber on a planned work piece with adaptive seals and partly immersed laser-head.1463Even though oxygen contents of about 4000 ppm in the protection-gas-atmosphere without significant impact on the mechanical properties have been reported, a threshold of 300 ppm has been chosen by experience in titanium welding. Applying this threshold has previously led to additively manufactured parts without discoloration from oxidation.

Applying this threshold has previously led to additively manufactured parts without discoloration from oxidation. To determine a minimal argon flow rate that provides the desired atmosphere in a feasible time, different flow rates have been tested. The oxygen content was measured by a zirconium based sensor with a relative error better than 5%. Through a hole in the base plate directly under the nozzle, the atmosphere was pumped to the sensor. The nozzle had a distance to the base plate of 6 mm, what is a typical process height. Since the atmosphere in the chamber might have deviations in the atmosphere composition at the beginning of each test, the data from 500 ppm to 300 ppm oxygen content was analyzed. The time from first reaching 500 ppm to the point where 300 ppm was not exceeded anymore was measured.

The time from first reaching 500 ppm to the point where 300 ppm was not exceeded anymore was measured. Figure 2 displays the graphs of these measurements with three different argon flow rates of 5, 6 and 7 liter per minute.Figure 2: Oxygen decay in protection gas atmosphere for different argon flow ratesThe first measurement, carried out at 5 l/min took 162 s to safely reach the desired atmosphere of 300 ppm oxygen content. With 6 l/min, the time was shortened to 15 s and with 7 l/min to 11 s. Since low argon flow rates are desired, 6 l/min argon flow rate was chosen for further testing.Controlling the oxygen content of the atmosphere is most important at elevated temperatures. With increasing distance to the processing area, the titanium had time to cool down and therefore the affinity to oxygen is reduced. This means more oxygen can be tolerated. A test has been conducted to measure the oxygen content in respect to the nozzle center, respectively the processing area.

This means more oxygen can be tolerated. A test has been conducted to measure the oxygen content in respect to the nozzle center, respectively the processing area. A value has been taken every other mm from a distance of 2 mm until 40 mm. Each measurement took roughly 3 min. The last 100 seconds of the measurement were used, taking a value every second. Therefore, 100 values were taken for each distance. The results are displayed in Figure 3. Within a distance of 36 mm to the center of the nozzle, the average oxygen content is below the threshold of 300 ppm. Away from the nozzle center, the standard deviation rises as the values increase. Within the 100 values taken for each measurement the maximum value was determined and never reached 300 ppm within a distance of 30 mm.1464Figure 3: Oxygen content in respect to the distance to the center of the nozzleTo test the setup a weld with Ti-6Al-4V has been performed.

A simple bi-directional meander with four lines has been welded on a substrate. The travel speed was 1000 mm/min, wire-feed was 2 m/min and the laser power 1100 W. No discoloration from oxidation is visible on the specimen. Therefore, it is clear that the protection-atmosphere was intact. Figure 5 displays a microsection from the welded Ti-6Al-4V sample shown in Figure 4. The microsection has been grinded and polished and etched with “Weck”.Figure 4: Ti-6Al-4V weldA martensitic microstructure rapid cooling can be seen in the microsection. Alpha- casing due to oxygen in the protection gas atmosphere would be displayed as a white region on the border of the weld. It absence underlines the feasibility of the low argon gas flow welding technique.fromFigure 5: Microsection of the welded Ti-6Al-4V sample1465Conclusion and OutlookThe surface quality of the welded specimen which does not show any discoloration from oxidation proofs the feasibility of the protection concept.

With 6 l/min argon gas flow, it is possible to weld structures that comply with the needs of Titanium parts. The concept allows to build 1.16 g titanium per liter argon what is more than five times the mass per liter argon reported by [13][14]. A drawback is the limited welding length inside the chamber. This constrain can be limited by a smart welding strategy depicted in Figure 6. Patches, consisting of multiple layers can be welded with slopes on the sides where they have to be prolonged in order to form the desired structure. The slopes allow to connect a new patch to the previous patch. This strategy allows to move the chamber successively over the substrate and will be demonstrated in further investigations.Figure 6: Patch based weld strategyAcknowledgementThe project has been founded by NBank, Projekt Vertical E2E ID ZW1-80159727References[1] Guleryuz, H.; Cimenoglu, H.: Oxidation of Ti-6Al-4V Alloy: J. Alloys Compd. 2009, 472, 241–246.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceA Solid Free Form Fabrication Equipment to Manufacture Axisymmetric Parts with Improved Surface QualityRajeev Dwivedi, Indira Dwivedi, Arihant Panwar, Bharat DwivediAbstractCompetitive and Hobby grade Rocket makers quite often build custom nozzles. Solid freeform fabrication is most natural choice for Manufacturing of the Nozzles. Different geometries can be quickly manufactured and tested. However, staircase effect and limited accuracy of 2-1/2 based deposition prevents the design intent from fabrication. Additionally, using different blends of ceramic and sustaining the geometry during curing becomes challenging. This research presents a unique 3D printing system that dispenses ceramic to enable manufacturing of axi-symmetric parts as continuous bead.

This research presents a unique 3D printing system that dispenses ceramic to enable manufacturing of axi-symmetric parts as continuous bead. Relative motion of the material dispenser and rotational substrate as well as unique path planning enables a continually sculpted surface to reduce the staircase effects.IntroductionHobby grade Rocketry has recently become very popular with learners, educators and (Do It Yourself) DIYrs. These rockets may range in complexity and size. In addition to the design for aerodynamics the rocket builders’ experiment with the controls, rocket fuel, rocket motor design, materials for rocket and motors etc. Different aspects of Robotics become natural vehicles for experimentation in science, technology, engineering, and mathematics (STEM) concepts. Additionally, many learners participate in various rocketry competitions. Mitchell et al.

Additionally, many learners participate in various rocketry competitions. Mitchell et al. [1] implemented a 10-week project-based course in rocket building project that combines theoretical content from several subjects with a high-order learning approach (create, evaluate, analyze) to advance the engineering skills of students. Eugene [2] suggests a program on practical rocketry that introduces young students and their teachers to the aerospace environment-including the design, build, and fly process; teamwork; and competition. The American Rocketry Challenge (ARC) is an aerospace design and engineering event for teams of US secondary school students (6th through 12th grades) run by the NAR and the Aerospace Industries Association (AIA) [3]. TARC is one of the most popular high school rocketry competitions in the United States. Teams of student’s design, build, and launch rockets to meet specific altitude and time duration goals.

TARC is one of the most popular high school rocketry competitions in the United States. Teams of student’s design, build, and launch rockets to meet specific altitude and time duration goals. The competition often involves restrictions on rocket size and motor type, fostering creativity and problem- solving. NASA Student Launch [4] is another research-based, competitive, experiential exploration program that provides cost-effective research and development of rocket propulsion systems for middle and high schools, colleges, and universities across the nation. Teams design and launch rockets with scientific payloads. Teams must adhere to specific guidelines and safety regulations.While primarily a college-level competition, some high school teams may participate in (Intercollegiate Rocket Engineering Competition) IREC.

It challenges teams to design and launch high-powered rockets with specific objectives, such as reaching a particular altitude or deploying a payload.In addition to these competitions, many regions have local rocketry competitions for high school students, organized by schools, community clubs, or educational institutions with varying rules and objectives. As a part of learning and experimentation, many teams design and manufacture their own1014rockets. With the focus on altitude gain and accuracy, many participants choose to design their own rocket motors.The research presented in this paper is based on developing nozzles for hobby grade rockets. The authors intended to design, manufacture and then experiment with different geometries popular within the hobby grade of rocket motors. To manufacture motors for experiment, 3D printing was a natural choice.

To manufacture motors for experiment, 3D printing was a natural choice. Availability of cost-effective 3D printer to manufacture heat resistant ceramic materials, cermits and other alloys was a limitation. Additionally, standard off the shelf as well as open-source 3D printing platforms use STL file, a 2-1/2 axis spatial manufacturing with zigzag path to deposit materials in sequential layers. Each being a challenge, authors pivoted to developing a novel 3D printing method amenable to axi-symmetric parts.Over the course of development, availability of both propellants as well as ceramic material became a limiting factor to build actual propulsion units. However, authors realized that the unique manufacturing method developed by them had many novel and unique features. The research shifts from traditional cartesian system-based approach to a polar coordinate-based approach for representation and manufacturing.

The research shifts from traditional cartesian system-based approach to a polar coordinate-based approach for representation and manufacturing. The paper presents the details of (1) 3D printer architecture (2) part and process representation (3) input file for the electrical actuator drives (4) extrusion and delivery of the ceramic materials. We present the preliminary results and some of the future developments based on findings.Modeling Axisymmetric partsTraditional additive and subtractive manufacturing processes are based on the geometric features of the parts being manufactured. For example, axi-symmetric parts may be manufactured using a lathe. By adjusting the rotational speed, pitch and tools, the process is optimized for rough machining and then fine machining. Similarly welding based repair of axisymmetric parts will benefit by adding the material while substrate is rotated about its axis.

Similarly welding based repair of axisymmetric parts will benefit by adding the material while substrate is rotated about its axis. On the contrary, the 3D printers due to 2-1/2 axis architecture is limited to depositing the material that may not necessarily align with the geometry (Figure 1). The material is deposited such that the end effector follows some variant of zigzag path. Additionally, the part will be susceptible to staircase effect (Figure 1). If part features and surface smoothness is critical, the surface is smoothened by post-processing.Figure 1 : Staircase effect and1015The work reported in this research suggests an alternate architecture. The architecture is limited to axisymmetric parts.6-axis robotic systems have been used to fabricate various axisymmetric and similar parts[5]. Yaou et.

The architecture is limited to axisymmetric parts.6-axis robotic systems have been used to fabricate various axisymmetric and similar parts[5]. Yaou et. al [6] reported a robot controlled LBDMD system that couples a 6-axis robot arm with an additional 2- axis tilt and rotatory positioning system to manufacture complex revolved parts. Additionally, they coupled 2-axis tilt and rotatory system and a hybrid slicing method to map the overhanging structures of a revolved part. Ramos et. al [7] reported fabrication of metallic solid objects using an experimental spiral growth selective laser melting (SG-SLM) 3D printer. The system incorporates a cylindrical coordinate system, instead of the ubiquitous Cartesian coordinate system found in commercial powder bed fusion selective laser sintering/melting (SLS/SLM) equipment.

The suggested systems above use expensive multi-DOF robotic manipulator.The process to manufacture suggested in this work focuses on creating a continuous path for material deposition. Figure 2 describes an improved approach where material is deposited as circular contours [8]. However, when transitioning from one layer to another, dwelling is introduced. Surface is not smooth along the dwelling point. Excess material added during the transition may be controlled by limiting the material flow. However accurate estimation of material and other physical properties of the material such as surface tension will impact the surface quality. Deposition processes that are based on heat energy (lasers, welding, polymer melting) will see higher energy input during the dwell, hence cause geometric as well as microstructural variations.Figure 2 : axisymmetric part with contours and dwell at layer transitionAs described in Figure 3, we are using a continuous path to overcome the dwell.

The material flow is uniform across the path; however, the speed of the end-effector is adjusted to maintain uniformity.Figure 3 : Schematic of continuous pathSystem descriptionFigure 4 Describes the system used for 3D printing. The system comprises of 2 linear actuators, one rotary actuator and material dispenser. Vertical linear actuator (Y) traverses in vertical direction at a constant speed during the material deposition process. Second linear actuator (X) actuates the material deposition end-effector in horizontal direction to adjust the radial distance.1016Figure 4 : 3D Printer SchematicThe rotary table (R) spins the substrate. The linear actuator uses a Misumui linear precision guide driven by haydon-kirk E46443-05-028ENG linear actuators. The rotary table is actuated using a general purpose NEMA23 stepper motor.

The linear actuator uses a Misumui linear precision guide driven by haydon-kirk E46443-05-028ENG linear actuators. The rotary table is actuated using a general purpose NEMA23 stepper motor. The system is controlled with open-source Arduino stepper motor module.Figure 5 : Modeling the nozzle geometry based on geometric simplification and continuous pathCoding for the system was done per part basis. Figure 5 describes a sample axisymmetric geometry. Part is divided into multiple segments and then a continuous spiral is modeled. The geometric coordinates of the spiral are identified and converted to corresponding stepper motor speeds. Given the scope of this project and similarity of nozzle geometries we wrote a custom speed-based program.

Given the scope of this project and similarity of nozzle geometries we wrote a custom speed-based program. Authors intend to write a general-purpose software that enables usage of open-source libraries to create GCodes in future.Figure 6 : Code implementation schematic1017As described in Figure 6, to enable uniform flow of material, the vertical speed is constant. Material is fed at uniform rate. Adjustment between the Radial and rotational speed is done in such a way that the rate of material dispense is constant throughout the manufacturing.Material and material dispensing SystemHigh performance ceramics, such as Ceramic Matrix Composite (CMC) are used for liquid rocket engines [9,10]. The materials are expensive and have limited availability. Additionally, the manufacturing process for the materials is specialized.

The materials are expensive and have limited availability. Additionally, the manufacturing process for the materials is specialized. As described in the earlier sections, the research work started with assumption of ease of material; however, the material could not be available therefore we pivoted to using an easily available ceramic Calcium sulfate hemihydrate (10034-76-1). For testing purposes slurry was created with 1-part water and 2-part Ceramic by volume.Figure 7 Universal Syringe adapterInitial attempts for material dispensing using an off-the-shelf Universal syringe adapter (Figure 7) were not very successful. The plunger is actuated pneumatically. Pneumatics could not be controlled to sustain uniform flow and material flow is non-uniform. The pneumatic system was replaced with an electronic linear actuator. Controlling the rate of actuation as well as the load is significantly improved with electronic actuation system.

The pneumatic system was replaced with an electronic linear actuator. Controlling the rate of actuation as well as the load is significantly improved with electronic actuation system. The system comprises of a screw that is actuated using stepper motor.Figure 8 : Stepper Motor actuated material dispense unitThe material dispensing unit utilizes a 10mL Luer Lock Syringe as reservoir and dispenser. The piston for the syringe is actuated using a Haydonkerk Size 8 Series 21000 Non-Captive Linear Actuator (Figure 8). The actuator is 3.56mm screw with 0.04mm lead per step.1018Initial ResultsFigure 9 : CAD model of the tested partFigure 9 describes the geometry that was tested for axisymmetric manufacturing. Part is 50mm tall. Two set of syringes were tested for material dispensing.Figure 10 : manufactured partsPass 1 Implementation: First experiment for dispensing was performed with 16 Gauge needle (1.6mm).

Two set of syringes were tested for material dispensing.Figure 10 : manufactured partsPass 1 Implementation: First experiment for dispensing was performed with 16 Gauge needle (1.6mm). The slurry starts to run and therefore frequent interventions are needed for the material to retain the shape. With the build up of the semi-solid liquid, at approximately 5mm thickness, the material structure starts to sag and part is not able to retain its shape as intended.Pass 2 Implementation: Second experiment is performed with 18 Gauge needle (1.2mm). We see a reduction in the slurry run. However, settling of the particles within the material reservoir causes intermittent obstructions to material flow; therefore, frequent interventions are required. We also adjusted the ratio of water in the slurry to 1-part water to 2.25-part ceramic.

We also adjusted the ratio of water in the slurry to 1-part water to 2.25-part ceramic. The uncured material is structurally stable; however, at approximately 7mm thickness material starts to deform.Pass 3 Implementation: Third experiment is based on improvements observed in the phase 2. Material is dispensed using 18 Gauge needle (1.2mm). The slurry has 1-part water to 2.25-part ceramic. The settling of material in the reservoir is reduced by tapping at interval of approximately 2 seconds. The ceramic is partly cured on the top surface by blowing hot air using 1200W hair dryer. The Slurry run is minimized and the instances of material flow obstructions are reduced.Conclusions and Future workA 3D printer and process that can be used for printing axi-symmetric parts was demonstrated. Part specific code was implemented.

Part specific code was implemented. The system was implemented using open source control hardware.1019Process was improved by (1) adjusting the ceramic-water ratio (2) preventing the ceramic particles from settling (3) curing the ceramic along surface using hot air. The development was motivated with intent to developing nozzles for hobby grade rockets; however, timely material availability limits the developments. The researchers will improve the printer by adding modules that prevent ceramic material from settling. Additionally, researchers would create an code to convert any axi-symmetric geometry to a G-Code for implementation with open-source hardware.References:.1. R. Mitchell Spearrin, Fabio A. Bendana, "Design-build-launch: a hybrid project-based laboratory course for aerospace engineering education" , Acta Astronautica, Volume 157, 2019, Pages 29- 39,ISSN 0094-57652. Fleeman, Eugene.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA SPHERICAL TEST ARTIFACT TO EVALUATE THREE-DIMENSIONAL FORM ACCURACY FOR WIRE ARC ADDITIVE MANUFACTURINGS. Ko*†, T. Sagawa*, Y. Yamagata*, S. Aoki*, and T. Abe**Institute of Technology, Shimizu Corporation, Japan †Corresponding author, E-mail: s_ko@shimz.co.jp **Division of Mechanical Engineering and Science, Saitama University, JapanAbstractAdditive manufacturing, including the wire arc additive manufacturing (WAAM), is gradually gaining attraction, and providing benefits in the aerospace and construction industries. In both industries, large-scale manufacturing capability and quality consistency of manufactured 3D parts are crucial. As part of quality evaluation, test artifacts for the geometric capability assessment are specified in ISO/ASTM52902-2019(E).

As part of quality evaluation, test artifacts for the geometric capability assessment are specified in ISO/ASTM52902-2019(E). On the other hand, the test artifact for curved wall is left undefined. This paper proposes a spherical shell shape as a representative of three-dimensional shapes that are supportless and feature large overhangs, for testing the geometric capability of a WAAM equipment. A mechanical configuration and deposition strategy are considered, which owns the potential to universally applying for depositing large-scale parts. A quality evaluation process for the sphere deposition was also described and experimentally demonstrated.Keywords: Wire arc additive manufacturing (WAAM), Test artifact, Inspection process, SphereIntroductionThe combination of an electric arc as heat source and wire as feedstock is referred as Wire Arc Additive Manufacturing (WAAM) [1, 2].

The combined use of multi-axis robot systems and wire arc additive manufacturing (WAAM) offers tremendous advantages in the flexibility and the large- scale manufacturing capability [3]. The same with other metal AM technologies, WAAM can save manufacturing lead time [3], reduce the waste, therefore lower the environment impact [4,5]. WAAM process can also meet the emerging flexibility requirements, such as integrated printing of complex shapes through functional optimization [5]. Thanks to the wire feed stock, the deposition rate of the WAAM process is basically 4 to 9 times compared to other AM technologies using powder by referring [5]. Generally, metal AM faces a limited surface finish level which deeply depends on the particle size of the powder or the diameter of the wire.

Generally, metal AM faces a limited surface finish level which deeply depends on the particle size of the powder or the diameter of the wire. Therefore, one of the typical applications of a DED (Directed energy deposition) technologies (including WAAM) is for the production that can accept a near net shape surface finish or with a subsequent machining process [6]. Although the discussion on the geometric accuracy of the wire arc additive manufactured parts could be found in [5, 16], the standardization on the evaluation and the assessment of the geometric errors by the WAAM process remains limited in the literature. For the additive manufacturing area, test artifacts for the geometric capability are proposed in ISO /ASTM52902-2019(E) [7]. However, only simple shapes such as hole, cylinder, slot, and straight walls are included. With the combination of the typical shapes, it is easy to access and possible to evaluate the basic geometric capability of the equipment.

With the combination of the typical shapes, it is easy to access and possible to evaluate the basic geometric capability of the equipment. On the other hand, curved1468wall, which is defined as a three-dimensional shell with different cross section in deposition direction [8] is difficult to be evaluated by ISO /ASTM52902-2019(E) [7]. Curved wall with large overhang feature can be deposited without printing extra support structures, by changing the direction of material accumulation during the fabrication process [3, 5], which is seen as a big advantage of WAAM process compared to other metal AM such as powder bed fusion (PBF) or direct energy deposition (DED) using metallic powder [8~11]. Therefore, the evaluation of the curved wall geometric capability is important for WAAM equipment.

Therefore, the evaluation of the curved wall geometric capability is important for WAAM equipment. To standardize the evaluation of the curved wall geometric capability, especially targeting to large scale workpieces, a sphere shape wall as a representative of three-dimensional shapes that are supportless and feature large overhangs, is proposed. A mechanical configuration and deposition strategy are considered, which owns the potential to universally applying for depositing large scale parts. A quality evaluation process for the sphere deposition was also described and experimentally demonstrated.A spherical test artifact for curved wall deposition evaluationWAAM is suitable for medium to large scale parts with low to medium geometric complexity [5]. The authors categorize the features available with WAAM process by the geometric complexity into Table 1. In which, the detailed feature of the Profile - Step 3 (multiple branches) can refer to [12].

The authors categorize the features available with WAAM process by the geometric complexity into Table 1. In which, the detailed feature of the Profile - Step 3 (multiple branches) can refer to [12]. The detailed feature of the Internal structure - Step 3 (lattice) can refer to [4, 13]. A spherical shell is a typical 3D shape that fully demonstrate the advantage of WAAM technology. In this study, the spherical shell is selected as the test artifact to evaluate the geometric accessibility of the WAAM equipment. Notice that, the deposition of multiple branches is more related to the path planning strategy, which is basically deployed by software development. The authors consider that the evaluation of sphere deposition can partially represent the multiple branches as well. On the other hand, lattice structure is not in the evaluation range of the proposed spherical test artifact.

Considering that the strong point of WAAM process is in the large-scale parts with geometric complexity up to medium level, lattice structure (high geometric complexity) with WAAM process is still far away from standardization and needs further research effort for the deposition feasibility and process stability control.Table 1 Available deposition objects with WAAM process Step 1: 2.5D shapeProfileStep 2: 3D shape with single branchStep 3: 3D shape with multiple branchesInternal structureStep 1: Shell modelStep 2: Solid objectStep 3: Lattice structureBy clarify the deposition strategy and geometric evaluation of the spherical test artifact with WAAM process, it would contribute to fabricate most of the target parts with curved wall feature.1469Strategy to deposit a large-overhang and support-free 3D shapeFor large-overhang objects deposition, extra support structure is usually necessary.

Some application with power-bed fusion [14] could deposit wall with a 10 ﾟ tilt angle from the horizontal plane. While for WAAM, or the traditional welding process, 30 ﾟ would be the limitation from our experience if we set the torch to be perpendicular to the horizontal plane. To overcome this limitation, a tilting rotary table with a swiveling axis was used to change the pose of the target part in [9, 10]. However, deposition of the parts in meters scale, which is generally used in the construction and the aerospace industries, would be difficult if the tilting of the deposited part becomes necessary, since of the gravity effect. A rigid fixture and safety measures becomes necessary to tilt the meter scale part and consequently the production becomes high-priced. Therefore, we propose a deposition strategy that basically always keep the deposited part in the horizontal plane.

Therefore, we propose a deposition strategy that basically always keep the deposited part in the horizontal plane. The raw material, which is the wire (and the torch) is kept in the tangential direction with respect to the deposited base. As well, the configuration to fit the deposition strategy, is also considered. Notice that Relativity Space [11] accepted similar deposition strategy in the Stargate Series 3D printers, while limited research results are published in the academia. Curved wall [8], as described previously, is defined as a three-dimensional shell with different cross section in deposition direction. Fig. 1 shows the conceptual scheme of the deposition strategy for the curved wall. In Fig. 1, the center, and the rotation with an angular speed ω around Z-axis can be defined. The slicing surface will slice the 3D model layer by layer in the Z-direction. The pose of the tool (a torch shown in Fig.

The slicing surface will slice the 3D model layer by layer in the Z-direction. The pose of the tool (a torch shown in Fig. 1) will be in a tangential direction from previous layer to the depositing layer. The distance between these two layers (called layer thickness) is setup to be in the tangential direction of the contour. Since the wire feed speed [15] in a WAAM process is set to be constant basically, the angular speed ω (rad/min) need to be specified following linear travel speed (mm/min) F = Rω. In which, R (mm) represents the distance between the contour and the Z-axis, as shown in Fig. 1.Fig. 1. Conceptual scheme to deposit curved wall1470A schematic design of the WAAM equipment fit for the meters-scale parts commonly used in the construction and the aerospace industries, is illustrated in Fig. 2. It includes a 6-DOF (degree of freedom) industrial robot with a welding torch as the end effector, and a rotary table that rotates around Z-axis in horizontal plane.

2. It includes a 6-DOF (degree of freedom) industrial robot with a welding torch as the end effector, and a rotary table that rotates around Z-axis in horizontal plane. A metal wire will be feeding from the center of the torch during the WAAM process, depositing layers of metal on top of each other to a desired 3D shape. The motion generated by the slicing algorithm, will be transformed to the rotation of the rotary table and the motion of the industrial robot. The pose of the welding torch should be tilted in YZ-plane (around X-axis), to keep the torch in the tangential direction relative to the deposited part. Consequently, the translational movement of the industrial robot can be limited to the minimum in this configuration. The configuration can provide an excellent deposition envelope with a limited robot size. By limiting the motion of the 6-DOF industrial robot, the influence of the volumetric errors by the robot, can be limited to minimum [17].

By limiting the motion of the 6-DOF industrial robot, the influence of the volumetric errors by the robot, can be limited to minimum [17]. Moreover, it fits better to the deposition of a meters-scale part with large-overhang, compared to the process mentioned in [9], since it is basically impossible to swiveling the parts considering the height of the parts.Fig. 2. The configuration of the WAAM equipment with a 6-DOF industrial robot and a rotary tableTable 2 The Experimental configuration of the WAAM equipmentIndustrial robot: Yaskawa AR1440/YRC1000Rotary table:Yaskawa MOTOPOS-D250FWelding power: Fronius TPS400iTorch:Fronius MTB400i 22 ﾟFixture (chuck): NOBEL WP-300To verify the feasibility of the considered deposition strategy, a sphere with a radius of 100mm was printed. The equipment used is listed in Table 2. The deposition conditions are shown in1471Table 3. The deposited sphere is shown in Fig. 3(a).

The equipment used is listed in Table 2. The deposition conditions are shown in1471Table 3. The deposited sphere is shown in Fig. 3(a). Notice that Yaskawa MOTOPOS-D250F consists of one swiveling axis around X-direction and one rotary axis around Z-direction. In the sphere deposition process, the swiveling axis was not used and was keep in 0 ﾟ to make the top of the rotary table in the horizontal plane. As the base for the sphere deposition, a base cylinder with a diameter of 60 mm, was clamped by a manual screw chuck and aligned to the center of the rotary table.

As the base for the sphere deposition, a base cylinder with a diameter of 60 mm, was clamped by a manual screw chuck and aligned to the center of the rotary table. The deposition of the sphere started from the outer diameter of the base cylinder (60 mm) and ended in the diameter of 60 mm before the peak of the sphere, to prevent interference between the torch and the deposited part.ParameterTable 3 Deposition conditions ValueBase cylinder (A5052(JIS))D60mm H150mmWire materialA5356-WY(JIS)Wire diameter1.2 mmWire feed speed2.9 m/minCurrent41 AVoltage11.1 VTravel speed600 mm/minDeposition rate0.5 kg/hTarget height per layer1.95mmEnvironment Temperature24℃Inspection process and case study to evaluate the quality of the deposited 3D shape(a) Deposited sphere(b) Solid model(c) Analysis method(d) AnalysisFig. 3.

3. The profile analysis of the solid model generated by CT scanning1472Table 4 Suggested measurement parameters for spherical test artifactParametersCommandMeasured by CT ScanningInner sphere radius (Ri±1σ)97.425 mm97.431 mm ± 1.561 mmOuter sphere radius (Ro±1σ)100.425 mm102.274 mm ± 1.287 mmWall thickness (t±1σ)3 mm4.9 mm ± 0.6 mmVoid volume ratio0.005%The deposited parts need to confront the weaknesses that is typically exist in the WAAM process, such as the integrity (the state of being fully dense), and the geometric accuracy [5, 15, 16]. Nevertheless, the deposition strategy considered here could have some concern on the integrity, referring to the traditional welding process [18]. Therefore, a scanning on the deposited part by using X ray computed tomography equipment (CT Scanner) was chosen in this study, since it has the advantage of non-destructively inspecting the porosity and the geometric accuracy [19].

Here below, an analysis process on the deposited sphere by using CT Scanning, is introduced. The analysis process includes three characterizations, 1) void volume analysis; 2) profile analysis; and 3) wall thickness analysis. As pointed out in [2], thermal influence is crucial in WAAM process. The heat accumulates and the profile of the weld bead can be varied as the deposition progresses [2]. Therefore, the characterizations depending on the deposition height, are of the authors’ interest. Since they can provide the information necessary to improve the large-scale deposition process. The experiment was conducted on a CT Scanner (XT H225ST by Nikon Solutions) [20]. VGSTUDIO (by Volume Graphics) [21] is used for the CT data analysis. The 3D model generated by the CT Scanning are shown in Fig. 3(b). Especially for the 2) profile analysis and 3) wall thickness analysis, an analysis method on the 3D model generated by the CT Scanning is considered, as shown in Fig. 3(c-d).

3(b). Especially for the 2) profile analysis and 3) wall thickness analysis, an analysis method on the 3D model generated by the CT Scanning is considered, as shown in Fig. 3(c-d). The definition of the analysis is shown in Fig.3 (c). The reference point to align the 3D model of the deposited sphere with the designed sphere, is setup to be the starting layer on the base cylinder (diameter 60 mm). The center of the sphere O is calculated based on the measured radius of the sphere. The radius of the outer profile and the radius of the inner profile can be calculated. They are summarized depending on the deposition height (see Fig. 3(c)). By calculating the distance between these two sets of data, the wall thickness depending on the deposition height can be calculated. Notice that a software named Geomagic Wrap [22] is used to repair the solid model generated by CT Scanning. After that, the calculation algorithm is implemented under Rhinoceros and Grasshopper.

Notice that a software named Geomagic Wrap [22] is used to repair the solid model generated by CT Scanning. After that, the calculation algorithm is implemented under Rhinoceros and Grasshopper. The suggested measurement parameters for the spherical test artifact are listed up in Table 4. The measurement results by CT Scanning are shown as well.1. Void volume analysisThe void can be detected and visualized, by the Porosity/Inclusion Analysis Module in VGSTUDIO [21]. The void volume, depending on the deposition height, is illustrated in Fig. 4. It shows that, only some dozens of voids within 0.1 mm3 were detected, although there are multiple voids of 6 mm3 max, especially in the deposition height 0 mm. The reason could come from a special pulse welding setting used only in the first layer to get a better penetration. the void volume ratio is 0.005% even including the deposition height 0 mm, which is considered not critical for this case study.

the void volume ratio is 0.005% even including the deposition height 0 mm, which is considered not critical for this case study. We can conclude that the change of the torch pose doesn’t affect the porosity of the1473curved wall. Remark that the voids occurred at deposition height 195 mm, which could be the influence from the cumulative error between the robot trajectory and the deposited part.Fig. 4. The void volume analysis of the solid model generated by CT scanning2. Profile analysisThe cross section of the deposited sphere in XZ plane measured by CT scanning, is shown in Fig.5. The profile and the wall thickness can be observed and discussed qualitatively. To further analyze the profile, the analysis described in Fig. 3 was conducted and the outer diameter profile (O.D.) and inner diameter profile (I.D.) are shown in Fig. 6. We can observe that the deposited sphere was relative open in both end (60 mm diameter on the bottom and the top).

6. We can observe that the deposited sphere was relative open in both end (60 mm diameter on the bottom and the top). Shrink is observed in the middle (or the equator 200 mm diameter), the same trend as in Fig. 5. This could be affected by the cooling after the WAAM process, since more material is deposited around 200 mm diameter, compared with both ends. From Fig. 6, we can also observe that the profile error is around 5 mm. Notice that the error bar shows the standard deviation (1σ) of the dimensions around Z-axis.1474Fig. 5. Cross section of the deposited sphere measured by CT scanning110)m m105( e r e h p s e h t100f o s u d a Ri959054595145195245Deposition Height (mm)Outer profileInner profileCommand to RobotFig. 6. The profile analysis of the deposited sphere depending on the deposition height3. Wall thickness analysisThe wall thickness, which is mentioned as the weld bead width in [2], is shown in Fig. 7.

The profile analysis of the deposited sphere depending on the deposition height3. Wall thickness analysisThe wall thickness, which is mentioned as the weld bead width in [2], is shown in Fig. 7. The thickness varies depending on the deposition height sharply, especially in the beginning, the middle, and the end of the deposition. The thermal effect and the pose error of the robot, or even the gravity effect on the melting pool, could be the critical factors. However, the mechanism needs to be clarified in the next step research.14757)m m6.5( e r e h p s e h t65.55f o h t d Wil l4.54a W3.53 -54595145195245Deposition Height (mm)Measured Wall WidthCommandFig. 7. The wall thickness of the deposited sphere depending on the deposition heightConclusion and DiscussionThis paper proposed a spherical test artifact and its quality evaluation process, for testing the geometric capability of the WAAM equipment.

A mechanical configuration and the deposition strategy for depositing the proposed spherical shell shape was also considered. Since the mentioned WAAM system owns the potential to universally applying to large scale parts deposition, the authors will focus on the system design, software deployment and feedback system development of the proposed deposition strategy in the next stage. Moreover, from the case study demonstrated in this study, the following research direction would be of interest in academia and needs further research efforts from the authors’ understanding. -The measurement and the improvement of both profile accuracy and the wall thickness stability. -The influence of the error factors from WAAM equipment on the form accuracy. -The non-destructive testing methods to replace CT scanning, especially for large scale parts.References[1] S. W. Williams, F. Martina, A. C. Addison, J. Ding, G. Pardal, and P.

S. Jalui1, T. J. Spurgeon1, E. R. Jacobs1, A. Chatterjee2, T. Stecko3, and G. P. Manogharan1*1 Department of Mechanical Engineering, The Pennsylvania State University, University Park, PA 16801 2 Department of Engineering Science and Mechanics, The Pennsylvania State University, University Park, PA 16801 3 The Center for Quantitative Imaging, The Pennsylvania Stata University, University Park, PA 16801Corresponding Author: gum53@psu.eduAbstractMetal additive manufacturing using Laser-Powder Bed Fusion (L-PBF) technique has enabled the metal manufacturing industry to use design tools with increased flexibility such as freeform internal channel geometries that benefit thermofluidic applications such as heat exchangers. A primary drawback of the L-PBF process is the as-built surface roughness, which is a critical factor in such surface-fluidic applications.

A primary drawback of the L-PBF process is the as-built surface roughness, which is a critical factor in such surface-fluidic applications. In addition, complex internal channel geometries cannot be post-processed through traditional finishing and polishing methods, and require advanced finishing processes such as Abrasive Flow Machining (AFM). In this original study, the effects of AM design including geometrical changes at the inlets, internal channel and wall thickness of thin features are experimentally studied on Ti64 L-PBF parts. A novel surface roughness inspection technique using micro-CT data is also presented. The internal channels with larger dimensions underwent 40% improvement in surface roughness with no statistically significant change in diameter whereas the channels with smaller dimensions and bends had a 38% improvement in surface roughness accompanied by a 6% increase in diameter.

While there was as much as 30% improvement in surface roughness values, the thin walls less than 0.4 mm in dimension were deformed under the AFM pressure after just 5 cycles.Keywords: Laser-Powder Bed Fusion, Additive Manufacturing, Surface Roughness, Abrasive Flow Machining, Micro-CT Scanning, Hybrid AM1646Figure 1: Graphical abstract1. IntroductionMetal additive manufacturing has enabled the fabrication of complex geometries that could not be produced using conventional techniques. Several industries including aerospace, defense, and heavy machinery benefit by leveraging this in applications such as heat exchangers, heat sinks, fuel passages, and conformal cooling channels. Organic, free-form designs allow for these components to contain superior functionalities when compared to their conventional counterparts due advantages such as partlight- weighting, and optimized topologies.

The global metal heat exchanger market was valued at $15.3 billion in 2019 and estimated to grow to $20.5 billion by 2024 [1]. Additionally, the injection molding industry currently stands at an impressive $60 billion in the United States alone [2]. As per the AM Power report 2020, the aerospace, turbines, and helicopter industries accounted for 29% of all AM system sales revenue [3].However, one of the biggest drawbacks of the metal Laser-based Powder Bed Fusion (L-PBF) AM process is the as-built surface roughness which typically ranges from 5-20 microns (Ra)[4][5][6][7]. Surface roughness is a critical factor for applications that include inaccessible features commonly found in gas turbines, heat exchangers, and injection molding conformal cooling channels. It impacts factors such as heat transfer, flow rate, boundary layer velocity, and ultimately efficiency of these components and their respective designs.

It impacts factors such as heat transfer, flow rate, boundary layer velocity, and ultimately efficiency of these components and their respective designs. Several finishing and polishing techniques have emerged such as milling, turning, abrasive jet machining, electro-chemical polishing, vibratory finishing, and abrasive flow machining, among others, that enable achieving the desired surface roughness values to ensure optimal performance. However, several challenges exist with using these processes on additively manufactured parts and their respective design intricacies. The critical surfaces within such internal passageways can be extremely hard1647to reach by the tooling that is required in some of these finishing and polishing methods. Abrasive flow machining (AFM) is one such method that involves a visco-elastic polymer loaded with abrasives to enter and flow through complex internal passageways.

Abrasive flow machining (AFM) is one such method that involves a visco-elastic polymer loaded with abrasives to enter and flow through complex internal passageways. Petare and Jain (2018) list the different variants of the AFM process that have emerged over the years [8]. Recent studies have shown the effects of traditional AFM on both conventionally [9][10][11] and additively manufactured surfaces. Bouland et. al (2019) proposed a model to predict material removal and surface roughness and found that there is a “river-bank” effect created at flat edges as a result of the AFM process [12]. Han et. al. (2020) performed AFM on additively manufactured through holes [13] and helical passageways [14] and reported an improvement in the areal roughness (Sa) values after 10 cycles. Duval-Chaneac et. al.

(2020) performed AFM on additively manufactured through holes [13] and helical passageways [14] and reported an improvement in the areal roughness (Sa) values after 10 cycles. Duval-Chaneac et. al. (2018) performed multiple AFM cycles on SLM manufactured coupons using 4 different media with varying viscosity and analyzed the roughness evolution behavior along with residual stress analysis [15]. Kum et. al (2020) proposed a model to predict material removal caused by AFM on external AM surfaces using a combination of CFD and Neural Networks [16]. Fabrication of detailed features with small dimensions such as thin walls are subjected to large machining forces, capable of causing permanent deformations that may be critical to the application and structural integrity of the component.

Deformation of thin walls and changes in dimensions coupled with defects such as keyhole porosity within the L-PBF process, could largely vary the actual performance of heat exchangers and gas turbine components from simulated results.Another potential problem that arises in the context of AM parts is the method of inspection. In order to ensure non-destructively measured dimensions and surface roughness values of organically shaped surfaces, micro-CT scans serve as a viable option. However certain limitations such as specimen size and voxel approximation persists while using this process. Several studies show the use the micro-CT data for roughness measurements of additively manufactured samples. Snyder et al.[17] talks about the capabilities of CT scan software, and their abilities to detect and evaluate resolutions up to 1/10th of the scans’ respective voxel size.

Snyder et al.[17] talks about the capabilities of CT scan software, and their abilities to detect and evaluate resolutions up to 1/10th of the scans’ respective voxel size. Moreover, data from scans of 35 μm voxel sizes resulted in a decrease in resolution via smoothing out some of the smaller features.Townsend et al. [18] discussed the factors impacting the accuracy of areal surface data obtained using micro-CT for roughness measurements. They also established how internal and external surface data has the ability to impact edge boundary approximations and variations for similar surface textures. This, in turn, affected the roughness measurements using CT. Kerckhofs et al. [19] and Pyka et al. [20] investigated the use of micro-CT data to evaluate linear mean profile roughness using 2D slices of the surfaces under consideration.

Kerckhofs et al. [19] and Pyka et al. [20] investigated the use of micro-CT data to evaluate linear mean profile roughness using 2D slices of the surfaces under consideration. No statistical difference was found for surfaces having roughness values between 5-30 μm, in the means of Pa (an unfiltered raw profile’s average roughness) values derived using contact profilometry and micro-CT data. However, surfaces that contained sub-micron roughness values also had CT data that considerably overvalued the Pa values when compared to those obtained using a contact profilometry method, even with a 1.5 μm voxel size.Carmignato et al.[21] investigated the effects of surface roughness on dimensional measurements obtained using CT data. The findings included lower roughness values calculated using micro-CT when compared to experimentally obtained values, caused due to voxel approximation of peaks and valleys.

The findings included lower roughness values calculated using micro-CT when compared to experimentally obtained values, caused due to voxel approximation of peaks and valleys. It was also reported that the standard deviations of such data increased across the board with an increase in voxel size, resulting from effects being more pronounced for larger voxel values.This paper focuses on studying the effects of the AFM process on the dimensions and roughness of additively manufactured titanium alloy specimens. The primary geometries under consideration were thin walls and internal channels with bends.16482. CAD designThe specimens were designed using Autodesk Fusion 360. Since the AFM process required fixturing, it was decided that the specimens will have a rectangular flange incorporated into the design to match the machining characteristics on our AFM machine. This ensured that the specimens could be swapped out easily.

This ensured that the specimens could be swapped out easily. Additionally, the specimens were designed as two halves that would ensure that they could be taken apart between AFM cycles to help with visual inspection and facilitate the cleaning process to remove any remaining media. A footprint of 50 mm x 45 mm was chosen for the rectangular flange. An area of 26 mm x 14 mm around the center was chosen for the housing of the channels. Four 9 mm diameter bolt holes were created on the corners of the rectangular flange. To bolt the two halves of the specimens, bolt holes for standard bolts with 5 mm diameter heads and 1 mm diameter threads were created. In order to make these self-supporting, they were designed in teardrop shapes with 55 degree angles. A 12 mm x 10 mm rectangle housed the actual geometry of the channels that varied across samples. The specimen designs were separated into three sets with each representing different features under consideration.

The specimen designs were separated into three sets with each representing different features under consideration. Set 1 consisted of a 10 mm diameter through hole. Set 2 consisted of rectangular slots with 1 mm, 0.5 mm, 0.25 mm, and 0.1 mm slot widths. Set 3 consisted of 4 mm diameter cylindrical channels with zero (through hole), one, and three bends. Each of the slots were separated by varying wall thicknesses as described in Table 1. To ensure that all surfaces were self-supporting, all bent, internal channel surfaces were at a 55 degree angle making the bend angle 110 degrees. A depiction of the cross-sections of all test specimen are shown in Figure 2.Table 1: Wall thicknesses for Set 2 specimensChannel Width (in mm)1 0.5 0.25 0.1Wall Thicknesses (in mm) 2, 1, 0.5, 0.25, 0.2 2, 1, 0.75, 0.6, 0.5, 0.4, 0.3 2, 1.5, 1, 0.75, 0.5, 0.4, 0.3, 0.25, 0.2, 0.1 2, 1.5, 1.25, 1, 0.75, 0.6, 0.5, 0.4, 0.3, 0.2, 0.1Figure 2: CAD drawings for all specimens.16493.

Experimental3.1. AM fabricationAM fabrication of Ti6Al4V specimens was performed using the 3DSystems ProX 320 L-PBF system. A 245W laser power, and 82 micron hatch spacing were used with a 60 micron layer thickness for a 100% dense part. The specimens were oriented such that the channels were printed in the z direction, perpendicular to the build plate.After fabrication, the samples were removed from the build plate using wire-EDM. Each half of the specimens were filed to ensure flat surfaces on mating faces. Tapping operations were performed on the bolt holes to ensure a secure fit between the 2 halves when bolted. No heat treatment was performed on the parts.3.2. AFM experimental procedureAFM was carried out using the AFM Vector machine by Extrude Hone. 2-way flow was used for this experiment.

No heat treatment was performed on the parts.3.2. AFM experimental procedureAFM was carried out using the AFM Vector machine by Extrude Hone. 2-way flow was used for this experiment. The media, commercially branded as EM 25103, which was the lowest viscosity media available, was chosen.The two halves of each specimen were bolted together and the entire specimen was bolted to the fixture placed between the 2 cylinders of the AFM machine. Once the media was loaded and the upper cylinder was lowered into position, the AFM parameters of 1300 psi (~9 MPa) media pressure and 1500 inch3 (~24.6 liters) were set. One AFM cycle was defined as one upward cylinder motion followed by one downward cylinder motion. 5 AFM cycles were performed on the Set 2 (rectangular channels and thin walls) specimens and 15 AFM cycles were performed on the Set 1 (10 mm diameter through hole) and the Set 3 (4 mm diameter bends) specimens.

Cycle times for each specimen were noted.Following AFM, Set 2’s thin walled channels contained excess media buildup, and were stripped of excess abrasive material that was left behind via using carrier media to connect to leftover media in the channels. The media’s respective surface tension was utilized, and in doing so, removed a majority of the material from the inlets and outlets of the channels, but proved to be insufficient towards the center of the specimen. Following the preliminary cleaning, the specimens were subjected to multiple cycles of ultrasonic cleaning in a bath of distilled water, rinsing out the broken-up media; drying off with compressed air. This cleaning cycle was repeated until no more abrasive media were found to depart from the small slits of the channels.3.3. Micro-CT scanning procedure & inspection methodologyConsidering the dimensions and geometry of the specimens, micro Computed Tomography (CT) was used as the method for inspection.

Micro-CT scanning procedure & inspection methodologyConsidering the dimensions and geometry of the specimens, micro Computed Tomography (CT) was used as the method for inspection. Micro-CT scans were obtained for as-built and post-AFM samples using the GE v|tome|x L300 multi-scale nano/micro-CT system available at the Centre for Quantitative Imaging (CQI) at The Pennsylvania State University. A voxel size of 12 microns was chosen to obtain a high resolution sufficient to extract surface roughness data while maintaining reasonable scanning times.The micro-CT data was obtained as .vol files which were imported to ImageJ for processing and measurements. After converting the data from 32-bit to 8-bit for ease of handling and computation, the slices were cropped to retain the data only containing the sample.

After converting the data from 32-bit to 8-bit for ease of handling and computation, the slices were cropped to retain the data only containing the sample. The 8-bit images were exported as .tiff files for further analysis.AFM involves material removal and one of the objectives of this research was to study to change in dimension of the additively manufactured specimen features. Dimension analysis was performed in ImageJ for both as-built and post-AFM specimens using the line tool followed by the measure command. After importing the 8-bit tiff stacks into ImageJ, it was observed that the images had a gradual shift in the grayscale values at the boundaries of the specimen features. In order to obtain a clear distinction between the material1650and air, thresholding was applied to ensure the specimen material was represented using white pixels (255 grayscale value) while the air was represented by black pixels (0 grayscale value).

Rotation and re-slicing operations were performed to obtain the 2D-slice view corresponding to the cross-section of the internal channels. For the rectangular channels (Set 2), the dimensions under consideration were the slot width and breadth, whereas radius was measured for both halves of the cylindrical channels (Set 1 & Set 3). Measurements locations are illustrated in Figure 3. Measurements were taken at equal intervals along the channel length based on 2D-slice numbers. For the cylindrical channels, radial measurements were taken at 45, 90, and 135 degrees as illustrated in Figure 3(a). Prior, to taking measurements, appropriate scaling was performed to convert the pixel value to mm using the voxel size of 12 microns. The line tool was used to create straight lines using the angle guide and dimension values were obtained using the Measure command. Roughness analysis using micro-CT data was performed to analyze the change in surface roughness values of these internal channels.

Roughness analysis using micro-CT data was performed to analyze the change in surface roughness values of these internal channels. A novel method was developed to measure linear surface roughness using micro-CT data using ImageJ and Wolfram’s Mathematica program. This is described in Figure 4. After thresholding in ImageJ, the 2D-images were re-sliced such that the resulting slice represented the surface of interest. These were obtained using a linear re-slice command for the rectangular channels (Set 1), and using a radial re-slice command for the cylindrical channels. Appropriate cropping, rotating, and combining operations were performed on the channels with bends to obtain a straight profile for the channel surfaces prior to using radial-re-slice. After obtaining the surface of interest, the raw surface profile data was obtained using the Plot Profile command in ImageJ.

After obtaining the surface of interest, the raw surface profile data was obtained using the Plot Profile command in ImageJ. This command provided a plot of the average grayscale value corresponding to every pixel along the length of the image. In order to convert the average grayscale value to mm, the following expression was used:ℎ𝑒𝑖𝑔ℎ𝑡	(𝑚𝑚) = (𝑣𝑜𝑥𝑒𝑙	𝑠𝑖𝑧𝑒	(𝑚𝑚/𝑝𝑖𝑥𝑒𝑙) × 𝑎𝑣𝑒𝑟𝑎𝑔𝑒	𝑔𝑟𝑎𝑦	𝑣𝑎𝑙𝑢𝑒	 × 	ℎ𝑒𝑖𝑔ℎ𝑡	(𝑝𝑖𝑥𝑒𝑙𝑠))/255Figure 3: Illustrations of measurements taken for (a) radius of cylindrical specimens (b) slot width & breadth for rectangular channelsThe resulting data was imported into Wolfram’s Mathematica and converted to a time-series data where the x-axis represented the length (in mm) and the y-axis represented the height (in mm). Before,1651calculating the roughness values, the surface form was removed from the raw profile using the Lowpassfilter command.

Before,1651calculating the roughness values, the surface form was removed from the raw profile using the Lowpassfilter command. A cutoff wavelength of 2.5 mm based on the ASME B46.1 standard [22] was used to filter out frequencies higher than the provided cutoff. This was performed using a Fast Fourier Transform (FFT) in the Lowpassfilter command. The roughness profile was calculated by subtracting this result from the raw profile. Using the Mean, Max, and Min commands, the average roughness (Ra), maximum peak height (Rp), and minimum valley depth (Rv) values were obtained. The total profile height (Rt) was calculated by subtracting the Rv values from the Rp values. Roughness calculations were performed at three locations of every surface of each half for both as-built and post-AFM specimens.In order to better visualize the areas affected by the process, in addition to surface texture change, the tiff stacks were imported into Avizo.

Segmentation was performed on the stacks to extract the data corresponding to the material. The segmented data was converted to a surface using the Generate Surface tool within Avizo without any smoothing to retain surface detail. The resulting surface consisted of a large number of triangles which were reduced to about 1/10th of their original size and exported as STL files for ease of handling. This was used for both surface texture visualization and to create deviation maps. Geomagic Control X by 3D Systems was used to create the deviation maps. The STL meshes were imported into Control X as measured data and compared to the STEP files of the specimen using the 3D Compare tool.Figure 4: Illustration of the roughness using micro-CT methodology4. Results & DiscussionThe micro-CT scans showed that the as-built 0.1 mm slots were fabricated with insufficient space for media to flow through and thus no AFM was performed on it.

Results & DiscussionThe micro-CT scans showed that the as-built 0.1 mm slots were fabricated with insufficient space for media to flow through and thus no AFM was performed on it. Using the decided AFM parameters along with the choice of media resulted in approximately 6 hours of cycle time for the 0.25 mm slots, and therefore eliminated from this study. The remaining specimens were analyzed for cycle times, dimension, and surface roughness variations.Most significant cycle time variations were recorded for the Set 2 1 mm and 0.5 mm specimens. The cycle times reduced with increasing number of cycles. For the 1 mm slots, the cycle time reduced from 69 seconds to165226 seconds, and for the 0.5 mm slots the cycle times reduced from 755 seconds to 172 seconds after 5 cycles, respectively. These were a result of increasing slot width with material being removed by the cutting action of the abrasive particles allowing more media to flow through it.4.1.

These were a result of increasing slot width with material being removed by the cutting action of the abrasive particles allowing more media to flow through it.4.1. Surface roughnessPost-AFM surface roughness measurements were compared to as-built for each specimen. Figure 5 shows the change in surface roughness values before and after the AFM process for the Set 1 specimen. Average Ra value reduced by 40.67%, from 10.66 µm to 6.32 µm, and the average Rt value reduced by 47.45%, from 97.1 µm to 51.03 µm after 15 cycles. Another finding was the reduction in variation of surface roughness measured by a 64.2% and 69.68% reduction in standard deviation of the Ra and Rt values respectively.

Another finding was the reduction in variation of surface roughness measured by a 64.2% and 69.68% reduction in standard deviation of the Ra and Rt values respectively. It can be concluded that for the 10 mm diameter through hole Set 1 specimen, there was a significant improvement of surface roughness in both smoothing (reduction in average Ra & Rt) as well as uniformity (reduction in standard deviations of Ra & Rt) after 15 cycles.Each surface of each slot was analyzed for surface roughness in the Set 2 specimens. Ra values measured before and after AFM are presented in Figure 6. The average Ra value for 0.5 mm slots reduced in a range between 10.18% to 31.29%. There was also a reduction in standard deviation for all surfaces that ranged from 38.82% to 86.73%. The average Ra values reduced for all surfaces of the 1 mm slots ranging from 5.55% to 19.81%.

There was also a reduction in standard deviation for all surfaces that ranged from 38.82% to 86.73%. The average Ra values reduced for all surfaces of the 1 mm slots ranging from 5.55% to 19.81%. However, surfaces 2, 3, 4, and 5 showed an increase in post-AFM Ra standard deviation by up to 3 times as the as-built Ra standard deviation. Ra standard deviation reductions of 8.72% to 50.1% was observed for all other surfaces. It can be concluded that after just 5 cycles, there was a significant improvement in surface roughness values in terms of both smoothing and uniformity of thin wall surfaces for 0.5 mm slots. However, for the larger 1 mm slots, there was reduction in Ra values observed, but certain surfaces with thinner walls showed an increased variability of the Ra value that suggested non-uniform surface roughness.

Another observation was that the magnitude of reduction of Ra value was greater for the smaller 0.5 mm slots as compared to the larger 1 mm slots for the same number of cycles.Figure 5: As-built v/s post-AFM comparison of Set 1 specimen1653Figure 6: As-built v/s post-AFM comparison of Set 2 specimensThe Set 3 specimens results are presented in Figure 7. It showed similar trends as Set 1 with reduction in both average Ra values as well as reduction in standard deviation of the Ra values. The average Ra values reduced by 20.4%, 38.34%, and 38.46% for the control, 1 bend and 3 bend specimens, The standard deviation of the Ra values reduced by 58.74%, 43.09%, and 75.19% for the control, 1 bend, and 3 bend specimens.Figure 7: As-built v/s post-AFM comparison of Set 3 specimens16544.2. DimensionsDimension measurements were taken on the as-built and post-AFM micro-CT scans of each specimen. As per Figure 8 the radius of the Set 1specimen remained largely unchanged.

DimensionsDimension measurements were taken on the as-built and post-AFM micro-CT scans of each specimen. As per Figure 8 the radius of the Set 1specimen remained largely unchanged. This was validated using a one-way ANOVA test which gave a p-value of 0.543, concluding that the means had larger variation within the conditions as compared to between the conditions. However, it was noted that there was a reduction in the standard deviation by 39.92% which was likely a result of peaks being washed down by the process causing a slight increase in radius at those locations.Figure 8: As-built v/s post-AFM radius comparison for Set 1 specimenDimensions in Set 2 underwent significant change. Figure 9 shows the first visual observation revealed deformation of three thin walls of 0.5 mm, 0.4 mm, and 0.3 mm within the 0.5 mm specimen. After removing the media, there seemed to be certain particles lodged into the channels.

After removing the media, there seemed to be certain particles lodged into the channels. The micro-CT scan settings did not capture the region containing the largest deformation due to a shorter field of view. However, upon scanning a sample of the media revealed large abrasive grains (~2.63 mm) as well as traces of dense material having a CT signature similar to that of metals as shown in Figure 10. Such contaminants could be a part of previous workpieces that the media has been passed through causing chips to be incorporated into it, or could be a part of the current workpiece. Few of such contaminants having dimensions larger that the slot width, caused the thin walls to deform at the inlets under the media pressure.Figure 9: Image of the bent thin wall in the Set 2 0.5 mm specimen1655Figure 10: Micro-CT scan of the abrasive mediaOverall, the slot width of every slot in the 0.5 mm specimen, other than Slot 1, increased.

The largestdeviation of 0.216 mm was observed in Slot 3. Slot 1 showed a reduction due to the thin wall bend.For the 1 mm slot width sample, similar observations were made. The thinner 0.25 mm and 0.2 mm walls were deformed while the other walls were not. Slot 2 underwent the largest change of 5.85% in average slot width. The maximum deviation was 0.147 mm. Slot 1 showed reduction in average slot width due to deformation of the thin wall, whereas all other slots showed an increase in average slot width. Additionally, the standard deviation for slot width increased in the post AFM samples for all slots. This indicated an increase in the variability of dimension along the length of the rectangular channels.

This indicated an increase in the variability of dimension along the length of the rectangular channels. The results for the 0.5 mm and 1 mm samples are summarized in Figure 11.Figure 11: As-built v/s post-AFM slot width comparison for Set 2 specimens (a) 1 mm; (b) 0.5 mm1656For the Set 3 specimens, about a 6% increase in the average value of the radius was observed for all 3 specimens after 15 cycles. The standard deviation for the Control and 3 Bend specimens increased by 3% and 37% respectively while it reduced by 13% for the 1 Bend specimen. The results are summarized in Figure 12. Thus, the variability in the radius along the length of the channel increased for the Control and 3 Bend specimens and reduced for the 1 Bend specimen after 15 cycles.Figure 12: As-built v/s post-AFM comparison of radius for Set 3 specimensIn order to visualize the regions that were affected by the process, deviation maps were created using reconstructed CT data.

Figure 13 shows the deviation maps of the as-built and post-AFM CT-reconstructions superimposed over the CAD STEP file. Shades of yellow, orange, and red represent the surfaces that extend above the CAD surface indicating the inherent surface roughness of the as-built workpiece. Deviation map of the post-AFM specimen shows blue areas that indicated the recessed surface caused as a result of the AFM process. The reconstructed models display the changes in surface texture caused due to the material removal action of the abrasive grain on the surfaces of the internal channels. Figure 14 clearly shows the markings created by the abrasive grains along the direction of flow of media inside the channel after 15 cycles.Figure 13: As-built (left) v/s post-AFM (right) comparison of deviation maps for Set 3 3 bend specimen1657Figure 14: As-built (left) v/s post-AFM (right) CT reconstructions of the Set 3 1 Bend specimen5.

ConclusionAbrasive flow machining is one of the few finishing and polishing methods that effectively enables post-processing of additively manufactured internal channels to improve surface roughness. 8 different specimens were manufactured using L-PBF to represent thin walls, through holes, and bends as internal channel geometries. The thin walled specimens were subjected to 5 AFM cycles while the through holes and bend specimens were subjected to 15 AFM cycles. Out of the initial 8 specimens, only 6 were machined since the specimens with 0.1 mm and 0.25 mm slots could not be machined due to improper fabrication and unreasonable AFM cycle time, respectively. Thus, media selection is an important aspect of the AFM process planning.Micro-CT was used as a non-destructive method to evaluate changes in dimensions and surface roughness as a result of the AFM process.

For the larger through hole specimen, there was a reduction in both average value and standard deviation of surface roughness while there was a reduction in variability of radius along the length of the channel. For the thin walls and rectangular channels, it was observed that walls thinner than 0.5 mm were deformed by the AFM media pressure. There was a reduction in surface roughness for all internal surfaces of the rectangular slots. However, it was noted that the resulting surfaces had varying values of surface roughness based on the respective as-built/initial surface roughness value. It can be concluded that for a specimen having multiple ports of entry, it is not certain that all internal channel surfaces will have uniform surface roughness values after the 5 AFM cycles. For the specimens with the bend(s) and the smaller diameter through hole it was observed that there was a reduction in surface roughness and an increase in radius after 15 AFM cycles.

For the specimens with the bend(s) and the smaller diameter through hole it was observed that there was a reduction in surface roughness and an increase in radius after 15 AFM cycles. A reduction in standard deviation of Ra values was observed but the standard deviation of radius of the 3 bend and through hole specimen increased after the AFM process. Overall, it can be concluded that internal channels having cross-sections of smaller dimensions (rectangular slot widths & 4 mm diameter cylindrical channels) showed increased variability in dimensions whereas the larger dimension specimen (10 mm diameter through hole) underwent no change in dimension while reducing variability along the length of the channel. It must be noted that all these conclusions are made for the specific media using which the experiments were performed. Future studies would incorporate a design of experiments with varying media while keeping the process parameters unchanged.

Future studies would incorporate a design of experiments with varying media while keeping the process parameters unchanged. More cycles1658could also be performed on these channels to assess if the process is able to reach uniform, desirable surface roughness values and the variation in dimension observed in the process.The AFM process resulted in the reduction of surface roughness of additively manufactured internal surfaces but caused deformation and dimensional changes. These variations are critical for the functionality of several potential applications that could benefit from the geometric complexity achieved using the L-PBF process. It is necessary to plan for these outcomes and create compensation models to incorporate them into the manufacturing process plan at the design phase itself.References[1]“Heat Exchangers Market worth $20.5 billion by 2024.” [Online]. Available: https://www.marketsandmarkets.com/PressReleases/heat-exchanger.asp.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAccelerated Corrosion Behavior of Additive Manufactured WE43 Magnesium AlloyRakeshkumar Karunakaran, Sam Ortgies, Ryan Green, William Barelman, Ian Kobler Michael Sealy1Mechanical and Materials Engineering Dept., University of Nebraska-Lincoln, Lincoln, NE 68588AbstractMagnesium alloys are capable of withstanding the high temperatures and pressures needed in oil and gas fracking operations followed by rapid and complete dissolution in days. Dissolvable magnesium plugs are used in fracking to enable longer lateral wellbores by eliminating mill-outs and the associated debris clogging. To increase extraction efficiency, the key technical challenge is determining how to increase the strength of a high corrosion rate magnesium device that enables higher pressures while maintaining high corrosion rates.

Topologically modified dissolvable plugs fabricated by additive manufacturing is proposed as a solution to fabricate high strength and high corrosion rate fracture plugs. Corrosion of magnesium is dependent on surface area exposed to corrosive media and is easily manipulated by additive manufacturing. This study highlights the development of optimal powder bed fusion process parameters for WE43 magnesium alloy and investigates the corrosion behavior of printed WE43 in a salt solution concentrated with sodium bicarbonate to initiate highly accelerated corrosion. Printed WE43 corroded three times faster than an as-rolled sample and was driven by the mechanical and materials properties formed by printing.Keywords: magnesium, corrosion, powder bed fusion, fracking1. IntroductionTraditional plugs used in hydraulic fracturing are composed of metal or composite materials.

IntroductionTraditional plugs used in hydraulic fracturing are composed of metal or composite materials. Fracking plugs isolate the lower zones of wellbores during multi-stage high-pressure fracking and are removed by milling after hydraulic stimulation is achieved. Milling of fracking plugs adds to production costs and elevate the risk of milling debris clogging wellbores. Traditional plugs also limit the lateral extension of borewells due to mill-out challenges in lengthy horizontal zones. As a result, there is a need for fully dissolvable fracking plugs [1]. The key barriers to efficient hydraulic fracking using dissolvable plugs are the poor mechanical properties and degradation performance of existing fracking plugs. Degradation rate of commercially dissolvable plugs are currently viable only for 2-mile lateral bores [2]. Dissolvable plugs do not exhibit high degradation rates reliably as required for shorter laterals of 1 to 1.5 miles and still need to be milled to remove.

Dissolvable plugs do not exhibit high degradation rates reliably as required for shorter laterals of 1 to 1.5 miles and still need to be milled to remove. Thus, better dissolvable fracking plugs are required for efficient fracking.Magnesium (Mg) and its alloys offer a moderate strength metal capable of degrading when in contact with water. Surface area directly correlates to the corrosion rate of magnesium and is a key parameter in developing high corrosion rate (HCR) devices. Additive manufacturing (AM) enables the use of more complex design tools, such as topology optimization, to fabricate1 Email sealy@unl.edu | Phone: 402-472-165962magnesium devices with varying degree of surface area exposed to corrosive environments.

Early investigations showed dense printed magnesium corroded, as measured by evolved H2 gas, at a rate of 5-7 mL/cm2 over 3 days in 0.1 M NaCl [3] compared to porous magnesium lattices that produced similar amounts of H2 in 27 days in 5 % fetal bovine serum [4]. These experiments were directed towards biomedical applications and used two different simulated body fluids (i.e., corrosion environments). To the authors’ knowledge, there is no literature on corrosion behavior of printed magnesium in highly corrosive environments. Additive manufacturing of magnesium plugs is hypothesized to address the HCR demands of hydraulic fracturing. The target dissolution rate is to have a fracking plug completely degrade in three days.Commercially available alloys for HCR dissolvable plugs include SoluMag® from Luxfer Mel Technologies and TAx-100E from Terves. The SoluMag® product is a variant of the WE43 alloy system.

The SoluMag® product is a variant of the WE43 alloy system. Thus, AM of WE43 magnesium alloy was conducted in this study to understand corrosion behavior. Salt water generated from fracking is three to six times more saline compared to seawater [5], and several sodium-based acids (accounting for 8 % of the solution) are supplied alongside fracking water to increase productivity [6]. Hence, HCR studies for printed WE43 were conducted in Hank’s balanced salt solution with 30X higher concentration of sodium bicarbonate to simulate oil fracking media. This publication adds to the knowledge on revolutionizing smart manufacturing of dissolvable, high-value magnesium tools for the oil and gas industry using three-dimensional printing of HCR devices.2. Optimizing Additive Manufacturing of Magnesium WE432.1 WE43 Powder and BaseplateWE43 magnesium powder, also known as MAP +43, was obtained from Luxfer Mel Technologies. The WE43 powder was spherical (Fig. 1a) with diameter 28 ± 7 µm (Fig.

The WE43 powder was spherical (Fig. 1a) with diameter 28 ± 7 µm (Fig. 1b). Parts were printed on precision ground (± 10 μm) 230 mm × 127 mm × 38 mm baseplates of magnesium alloy ZK-60a. The ZK-60a plates contains zinc and zirconium and is artificially aged at 149℃ for 24 hours. This baseplate is stronger and more ductile than WE43 and has a higher thermal conductivity (120 W/mK) compared to WE43 (50 W/mK).(a)(b)y c n e u q e r F14 12 10 8 6 4 2 010 15 20 25 30 35 40 45 50 Particle size (µm)Fig. 1 WE43 powder (a) morphology and (b) particle size distribution.632.2 Single Track Depositions of WE43Preliminary experiments were conducted to develop a suitable print recipe by depositing singles tracks using powder bed fusion (PBF) AM in a Matsuura Lumex Avance-25. The Lumex utilizes Yb-fiber laser with a maximum power output of 1000 W and a laser spot size of 0.30 mm. Single tracks 100 mm in length were deposited in an argon atmosphere with less than 1.5 % O2.

The Lumex utilizes Yb-fiber laser with a maximum power output of 1000 W and a laser spot size of 0.30 mm. Single tracks 100 mm in length were deposited in an argon atmosphere with less than 1.5 % O2. The PBF process parameters used for single tracks are listed in Table 1. Single track samples were sectioned off the baseplate and mechanically polished and etched with a 10% nitric acid solution for 5 seconds to reveal meltpool geometry. The meltpool was imaged with an AmScope polarizing metallurgical microscope at 5x magnification. The meltpool width and depth was analyzed using ImageJ software. Meltpool geometry attained from single track depositions was used to identify the optimal process parameter sets to fabricate cubes that yield the highest density. The layer thickness for depositions was 50 µm and the meltpool needed to penetrate approximately 100 µm to fuse adjacent layers.

The layer thickness for depositions was 50 µm and the meltpool needed to penetrate approximately 100 µm to fuse adjacent layers. The melt pool also needed to be 150 µm in width to allow for 50% overlap between adjacent rasters that are typically 100 µm in width [8].Table 1 Process parameters for single tracks on baseplate with and without WE43 powderProcess Parameters Laser Power: Scan Speed:100 W, 135 W, 200 W, 350 W 500 mm/s, 700 mm/s, 1000 mm/sA representative image of meltpool attained from single track depositions is shown in Fig. 2. The meltpools attained from single tracks were 100 μm ± 15 μm deep, 150 μm ± 20 μm wide, and showed symmetry about their center axis. No instances of key holing were observed in the depositions. The single tracks also adhered well with the baseplate.50 μmFig. 2 Meltpool obtained at 135 W laser power and 700 mm/s scan speed.Increase in laser power yielded larger melt pool while increasing scan speed reduced meltpool size.

2 Meltpool obtained at 135 W laser power and 700 mm/s scan speed.Increase in laser power yielded larger melt pool while increasing scan speed reduced meltpool size. Low laser powers and high scanning speeds lowered the interaction between the laser and the baseplate, leading to lower meltpool dimensions. The depth and width varied linearly with scan speed and had a cubic relation with laser power. A previous publication noted a parabolic relationship between laser power and width that may be attributed to differences in the PBF64system [9].

A previous publication noted a parabolic relationship between laser power and width that may be attributed to differences in the PBF64system [9]. Table 2 indicates four sets of single track process parameters that yielded optimal meltpools.Table 2 Meltpool dimensions for optimal process parametersLaser power (W) 100 135 200 200Scanning speed (mm/s) Meltpool depth (mm)700 700 500 7000.088 0.099 0.097 0.089Meltpool width (mm) 0.163 0.160 0.167 0.1652.3 Cubes for Density MeasurementsFour sets of laser power and scanning speeds were identified from single track depositions that yielded the ideal meltpool dimensions. The parameter sets coupled with three hatch spacings (0.08 mm, 0.10 mm, and 0.12 mm) formed 12 process parameter sets for cube depositions while the layer thickness was maintained at 50 µm. Three cubes were printed for each parameter set for statistical analysis of cube density.

Three cubes were printed for each parameter set for statistical analysis of cube density. Parts were printed from right to left on the baseplate to oppose argon gas flow in the system (Fig. 3). Average density for each set of three cubes was determined using Archimedes principle. Chemical composition and microstructure of the top three parameter sets that yielded highest densities were analyzed on a FEI Helios NanoLab 660.Printing direction for cubesDirection of Argon flowFig. 3 Cubes for density measurements printed on a baseplate.The optimal process parameter sets identified from single track depositions were used to print cubes by coupling with three different hatch spacings indicated in Table 3. All printed cubes were at least 96% dense compared to as-rolled WE43 that had a density of 1.845 g/cm3. Parts printed with laser energy density of 25 to 75 J/mm3 yielded highest density cubes.

Scanning electron microscopy of the highest density printed WE43 showed rose-like grains typical of printed WE43. Flake-shaped yttrium oxide was dispersed within the sample (Fig. 4).zyx xY-Z cross section analyzed by SEMYttrium oxide flake5 µmFig. 4 Microstructure of WE43 printed with laser power of 135 W, scan speed of 700 mm/s, and hatch spacing of 0.12 mm.3. Corrosion of Printed WE43An immersion corrosion test (ASTM G31-72 [7]) was used to examine dissolution of WE43. Magnesium reacts in aqueous environments to form magnesium hydroxide (Mg(OH)2) and hydrogen gas, H2. Magnesium reacts with water to form H2 in a 1:1 mole ratio that was collected and measured to examine corrosion rate of printed WE43 as a function of time. The corrosion66product Mg(OH)2 is basic and increased the pH of the solution. CO2 was circulated through the corrosion media intermittently to form carbonic acid and lower the solution pH.

The corrosion66product Mg(OH)2 is basic and increased the pH of the solution. CO2 was circulated through the corrosion media intermittently to form carbonic acid and lower the solution pH. The solution was renewed every 2 days and maintained between 7-8 pH throughout experiments.3.1 Corrosion Sample PreparationWE43 cylinders 38 mm in diameter and 7 mm in height were printed by PBF using optimal process parameters that yielded the highest density cubes (135 W laser power, 700 mm/s scanning speed, 0.12 mm hatch spacing, and 50 µm layer thickness). Printed samples were separated from the baseplate by a bandsaw, and the parted surface was milled to attain a final sample height of 3 mm. Both parting and milling operations were performed without coolant to avoid corrosion during sample preparation. A 1.5 mm hole was drilled at the center of samples to mount into the immersion setup. Samples were enclosed in epoxy in order to expose only the top surface to corrosion media (Fig. 5).

A 1.5 mm hole was drilled at the center of samples to mount into the immersion setup. Samples were enclosed in epoxy in order to expose only the top surface to corrosion media (Fig. 5). The exposed surface of samples was polished using 150, 320, and 400 grit sandpapers, cleaned with isopropyl alcohol (IPA), and dried. Density of all corrosion samples was measured by Archimedes’ method using IPA before immersion tests. Corrosion of printed WE43 was compared with commercially procured as-rolled WE43. The as-rolled samples were also mounted in epoxy and had the same surface preparation as the printed WE43 samples.PVC pipeEpoxy resinExposed surface of WE43 sampleFig. 5 WE43 sample for corrosion testing.Samples printed for corrosion studies were larger than samples for previous experiments in this publication. The quality of printed corrosion samples showed dependence on the location within the print chamber unlike prior WE43 prints.

The quality of printed corrosion samples showed dependence on the location within the print chamber unlike prior WE43 prints. Samples printed closer to print chamber door showed delamination compared to parts printed farther away (Fig. 6). Layer delamination was likely due to the directional flow of argon gas inside the print chamber. Soot generated during prints was carried by argon and excessively deposited around samples 1, 2, and 3. Soot has higher carbon content and is thought to have raised the melting temperature of the powder bed around the three samples. The 4th and 5th samples had minimal soot deposition and had higher heat retention while printing due to larger sample size. WE43 has a low melting magnesium alloy that led to WE43 powder vaporization within samples 4 and 5. Weak interlayer bonding due to powder vaporization is suspected to have caused layer delamination.

WE43 has a low melting magnesium alloy that led to WE43 powder vaporization within samples 4 and 5. Weak interlayer bonding due to powder vaporization is suspected to have caused layer delamination. Density of printed samples measured by Archimedes’ method showed samples 4 and 5 had 94 % relative density while samples 1, 2,67and 3 had 97 % compared to as-rolled WE43 (Fig. 7). No such delamination was observed for smaller cubes printed earlier to identify optimal process parameters.r o o d12r e b m a h c34Sample 1tn i rP5Sample 5Fig. 6 Printed WE43 corrosion test samples. Samples 1-3 showed no delamination while samples 4 and 5 delaminated.) 3m c / g (y t i s n e D1.85 1.83 1.81 1.79 1.77 1.75 1.73 1.71 1.69 1.67 1.65Wrought WE43% 8 1 . 7 9% 2 9 . 7 9% 8 2 . 6 9% 2 7 . 3 9% 4 4 . 4 912 4 3 Sample number5Fig. 7 Dependency of the density of printed WE43 corrosion samples on their position in the print chamber.

7 9% 2 9 . 7 9% 8 2 . 6 9% 2 7 . 3 9% 4 4 . 4 912 4 3 Sample number5Fig. 7 Dependency of the density of printed WE43 corrosion samples on their position in the print chamber. Percentage values represent relative density compared to as-rolled WE43.3.2 Immersion Testing – Hydrogen EvolutionA 210 mm × 450 mm × 210 mm glass tank was kept over heating plates for the immersion tests. The tank was filled with Hank’s balanced salt solution prepared by adding 10 L of deionized water alongside 97 g of dried Hank’s balanced salts (Sigma Aldrich: H1387-10X1L). The solution was doped with 100 g of NaHCO3 as opposed to the recommended 3.5 g to create a concentrated corrosion medium. The corrosion solution was renewed every 2 days and the temperature was set to 37 ºC. Two stir bars rotating at 200 RPM ensured uniform dispersion of salts and temperature in the solution during the immersion tests. Fig.

Two stir bars rotating at 200 RPM ensured uniform dispersion of salts and temperature in the solution during the immersion tests. Fig. 8 shows the setup for immersion experiments.68FlaskpH ProbeCO2 TankHank Solution Mg SampleHot Plate & MixerFig. 8 Experimental setup for corrosion testing.Samples were placed on a polycarbonate platform at the bottom of the tank, and a graduated beaker was placed above the samples to capture the evolved H2. A control beaker was placed inside the tank to account for gas bubbles formed due to cathodic hydrogen evolution reaction (reduction of water). Polypropylene hollow balls were added to the tank that floated over the solution to reduce evaporation of the corrosion media. The solution was allowed to heat for 6 to 8 hours prior to corrosion tests to remove dissolved gases before immersion tests. Solution temperature and pH were tracked regularly using a probe (ThermoFisher: 8107BNUMD).

Solution temperature and pH were tracked regularly using a probe (ThermoFisher: 8107BNUMD). CO2 was supplied at 1 psi when the solution media basicity exceeded 8 pH and was restored to 7 pH. The dispersion of CO2 was supplied by a diffusion tube with small pores to allow bubbles to form across the entire tank to maximize the interaction within the media. Evolved hydrogen was recorded every 15 minutes during experiments to track corrosion rate.Immersion experiments on printed WE43 indicated evolved H2 was highly sensitive to the pH of the corrosion media. Brittle Mg(OH)2 precipitates increased alkalinity of the solution that decelerated corrosion. Hydrogen evolved from printed WE43 accelerated when solution pH was lowered to 7 by suppling CO2 to the media, while H2 release plateaued when solution exceeded 7.5 pH (Fig. 9a).

Hydrogen evolved from printed WE43 accelerated when solution pH was lowered to 7 by suppling CO2 to the media, while H2 release plateaued when solution exceeded 7.5 pH (Fig. 9a). A pH dependent H2 release was also observed for as-rolled WE43, although fluctuations in H2 evolution due to pH was not as significant compared to printed WE43 (Fig. 9b). The solution alkalinity increased at 0.03 pH/hr. and led to the preferential formation of magnesium carbonate (MgCO3) precipitates at alkalinity higher than 8 pH. The carbonate precipitate formed a thick stable film over the samples and reduced anodic and cathodic interactions of the base metal with the corrosion media, effectively reducing hydrogen evolution from WE43 [10]. Pitting corrosion alongside porosity in printed WE43 increased exposed surface area to corrosive media that drastically accelerated corrosion in the latter portion of the immersion tests for both wrought and printed WE43.

Metal AM also accumulate tensile residual stress near surface that creates a favorable environment for accelerated pitting corrosion in printed WE43 [11]. Fluctuations in H2 evolution due to pH change in the solution smoothened out in the plot (Fig. 10) for corrosion rate. Overall, a 30-fold increase in NaHCO3 resulted a ten-fold increase in H2 evolution for printed WE43.69(a)70Printed WE43 Evolved H2pH(b)70As-rolled WE43 Evolved H2) 2m c /L m(2 H d e v o v El6050403020107.047.427.427.87.097.57.57.767.017.36) 2m c /L m(2 H d e v o v El6050403020107.297.577.578.017.097.667.947.047.5700010203040 Time (hr)506070010203040 Time (hr)506070Fig.

9 H2 evolution in concentrated NaHCO3 solution for (a) printed and (b) as-rolled WE43.500) y / m m( eta r n o s o r r o Ci400300200100Wrought WE43 As-rolled WE43Printed WE4300102040 30 Time (hr)506070Fig. 10 Corrosion rates of printed and as-rolled WE43 in concentrated NaHCO3 solution.4. Summary and Future WorkThe printability of WE43 magnesium alloy by powder bed fusion was investigated by varying laser power, scanning speed, and hatch spacing. Layer thickness was kept constant at 50 µm for all prints. Highest density cubes (1.843 g/cm3) were printed at 135 W laser power, 700 mm/s laser scan speed, and 0.012 mm hatch spacing. A relative density of 99.89% was attained compared to as-rolled WE43. A ten-times higher corrosion of printed WE43 was attained in Hank’s solution with 30-fold concentrated NaHCO3 compared to non-concentrated Hank’s solution.

A ten-times higher corrosion of printed WE43 was attained in Hank’s solution with 30-fold concentrated NaHCO3 compared to non-concentrated Hank’s solution. Pitting corrosion was accelerated by porosity and surface tensile stress due to additive manufacturing. H2 evolution from WE43 during corrosion accelerated at lower alkalinity (7-7.5 pH) and slowed at 7.5-8 pH due to the formation of stable magnesium carbonate films on sample surface. Powder bed fusion printed WE43 corroded three times faster (52.3 mm/year after 60 hours) compared to as-rolled WE43 (19.8 mm/year after 60 hours) in concentrated sodium bicarbonate solution. Future work is to perform accelerated corrosion experiments at higher temperatures to understand the dependence of the cathodic hydrogen evolution reaction on temperature within highly corrosive media.

Also, future work will analyze corrosion products to quantify the thickness of magnesium carbonate film on WE43 and its influence on corrosion rate.70AcknowledgementsThis research was conducted as part of an undergraduate/graduate investigative research course in additive manufacturing in the Department of Mechanical and Materials Engineering at UNL Spring of 2020. This activity was supported by NSF CMMI: 1846478 and performed in part in the Nebraska Nanoscale Facility (NSF ECCS: 1542182) and Nebraska Engineering Additive Technologies (NEAT) Lab supported by the Nebraska Research Initiative. Matsuura USA supported modifying the Lumex Avance-25 for printing magnesium.References[1] Walton, Z., Fripp, M., Porter, J., Vargus, G., 2019. Evolution of Frac Plug Technologies– Cast Iron to Composites to Dissolvable, SPE Middle East Oil and Gas Show and Conference, Manama, Bahrain.[2] Garcilazo Meza, G., Bautista Alarcon, X., Loundy, R., 2020.

Additive manufacturing of dense WE43 Mg alloy by laser powder bed fusion, Additive Manufacturing 33, p. 101123.[10] Johnston, S., Shi, Z., Atrens, A., 2015. The influence of pH on the corrosion rate of high- purity Mg, AZ91 and ZE41 in bicarbonate buffered Hanks' solution, Corrosion Science 101, p. 182-192.[11] Sealy, M.P., Karunakaran, R., Ortgies, S., Madireddy, G., Malshe, A.P., Rajurkar, K.P., 2021. Reducing corrosion of additive manufactured magnesium alloys by interlayer ultrasonic peening, CIRP Annals, in press.71.

tal'IJ''lUllUl'l."'I.,,U by1....~""l ....."ACh"'~ Institute for Applied Materials Research - IFAM,WilkeningEOSPlaneggmechanical properties of direct laser-sintered metal parts. and after an infiltration. Furthermore the accuracywere parts was parts show the microstructure of the copper-nickel-tin alloy.tested measured. Micrographs ofachievable complexity of parts is demonstrated by examples. An overviewactivities is/1/.metals was published last year describing a bronze process aswork shows the development ofdevelopment of the postprocessing by infiltration. Different increase surface prototypes as shown.have beenC'+~,,~.... n·th and hardness of the prototypes. Furthermore the ln11,,\rn'JPtYlPl"'1tby electrodeposition is shown.surface quality is necessarymolds.first polymer parts produced by injection 11r'\All"1u",nis a mixturenickel with copper, tin and phosphorus.are very common to by between 20 and 50 microns (Fig. 2). microns.

2). microns. sintering process.a low melting binder phase /2/.1 the sum analysis measured(energy-dispersive-analysis-of-X-rays) of the powder is shown.parts were built with a layer "&AA~"&"AAV"''''3 shows a(scanning-electron-microscope) image of sintering process works in ambient atmosphere at room "~A"''''''''''APOST-PROCESSINGpost-processing of the prototypes can be separatedtwo steps.infiltration to obtain dense parts and to achieve a first smoothing of the surface /3/. step is an electrochemical smoothing of the surface for the use as molds. Without smoothing the microscopic undercuts prevents part removal or cause the destruction of part during the removal. The infiltration was done with different infiltrants shown The infiltration processes was varied in time, atmosphere and use of fluxing agents ,...=....::.::;..;::;...;:... The best results could be achieved with the PbAg2Sn2 infiltrant. The specimen infiltrated with the other materials had a high porosity.

4) and Zamak were infiltrated with a water based fluxing agent (F SW12). The higher melting materials (Ag40Sn (Fig. 5), Ag40Cd (Fig. 6» were infiltrated with a pasty fluxing agent (F-SH1). The surface was smoothed by a nickel coating.Tab.2: Infiltration, experimentsZamak "30,60 30,60 5,45,90 5,45,90 30,60 30,60xxnone F-SW 12 F-SH 1 F-SH 1 none F-SW 12++ 0 + +MECHANICAL BEHAVIORThe mechanical behavior is shown in table 3. Last year we presented the first results using a lead based infiltrant. The result was an increase in tensile strength of about 60 % compared with the uninfiltrated sintered powder. The transition to silver based infiltrants led to much better results.

The result was an increase in tensile strength of about 60 % compared with the uninfiltrated sintered powder. The transition to silver based infiltrants led to much better results. The tensile strength could be increased from 132 MPa using the lead based infiltrant to 250 MPa by using a silver based one.Table 3: Mechanical behavior of the infiltrated specimenwithout infiltration L-A 40Sn L-A 40Cd PbA 2Sn281 250 130 13273ACCURACY AND APPLICATIONSOf the various production processes used for plastic components (extrusion, blow molding, etc.) injection molding is by far the most widely used for complex shaped parts of the type typically built by layer manufacturing techniques.A number of process chains and hybrid techniques have been used to produce injection molding toolings from stereolithography models, sometimes involving multiple stages of replication using both flexible silicone and rigid epoxy or similar castings.

The total achievable accuracy tends to decrease as a number of replication processes increases, and the number of prototypes which can be produced depends on both the final mold material and the material being injected, but hard tooling suitable for small series of prototype parts can be made.Hard tooling suitable for injection molding of a wider range of materials and larger series can be built directly in metal using EOSINT M.For dimensions up to 100 mm a building accuracy of approximately O. 1 mm can be achieved, for larger dimensions 0.1-0.2 %. The layer thickness is typically 0.1 mm. Figure 7 shows three stages in the production of a metal injection mold. On the left is a laser-sintered mold insert without any further post-processing. Next to it is the other half of the mold, after being infiltrated with tin.

On the left is a laser-sintered mold insert without any further post-processing. Next to it is the other half of the mold, after being infiltrated with tin. The infiltration fills in the pores between the sintered particles to produce almost full density and a much smoother surface, but has negligible effect on the part accuracy. On the right is an identical infiltrated mold insert after finishing the molding surfaces by hand. These finished molds were used for injection molding at the CRIF institute in Liege, Belgium. First the parting surfaces were machined by EDM, then the molds were embedded in a mother mold and runners and ejectors were machined. Subsequently several hundred parts were injection molded in polypropylene, polyethylen and ABS at a nozzle pressure of 600 bar and injection temperatures of up to 225°C. 35%-glass-filled nylon parts have also been injected at 280°C and up to 850 bar. After 300 shots with ABS there was no visible wear of the molding surfaces.

35%-glass-filled nylon parts have also been injected at 280°C and up to 850 bar. After 300 shots with ABS there was no visible wear of the molding surfaces. Figure 8 shows the mold and some parts produced by injection molding. Tests of the durability of the molds are continuing at EOS and a number of beta-sites.EOSINT M enables Rapid Prototyping directly in metal for the first time and thereby provides a much shorter process chain for technical prototypes in plastic than was previously possible. Further work is necessary to establish the technical and economic limits of these technologies, and intense development work and investigations of applications will continue for some time.FUTURE DEVELOPMENTIn future steel based powders will be used. First experiments have been done with coated 316L powders with good results.

First experiments have been done with coated 316L powders with good results. Powders with higher melting points and therefore other infiltrants with higher strength can be used.CONCLUSIONSParts directly laser-sintered from metal powders are suitable for many applications. It has been proved that the infiltrated material can be used as molds for injection molding. The strength of the infiltrated specimen has been increased from 132 MPa to 250 MPa by using other infiltrants74during post-processing. It has been shown that the results are sufficient for the production of prototype molds for injection molding.REFERENCES/1/ M. Sindel,Pintat, M. Greul, O. Nyrhila, C. Wilkening, "Direct Laser Sintering of Metals and Metal Melt Infiltration for Near Net Shape Fabrication of Components", Proc.oftheSymp. 1994, p.101/2/Bourell, Barlow, 1.W., Beaman, 1.1., "Selective Laser Sintering of Metals and Ceramics", International Journal of Powder Metallurgy (Oct.

5: Micrograph of the powder infiltrated with Ag40SnFig. 6: Micrograph of the powder infiltrated with Ag40Cd7850:150:1Fig. 7: Three stages in the production of a metal injection mold (left: laser-sintered, middle:infiltrated with tin, right: finished by hand)Fig. 8: Mold and the resulting molded parts79.

CURACY AND MECHANICAL PROPERTIES OF OPEN-CELL MICROSTRUCTURES FABRICATED BY SELECTIVE LASER SINTERING S. Eosoly*, G. Ryder†, T. Tansey†, L. Looney* * School of Mechanical and Manufacturing Engineering, Dublin City University, Dublin, Ireland †Department of Mechanical Engineering, Institute of Technology Tallaght, Dublin, Ireland Abstract This paper investigates the applicability of selective laser sintering (SLS) for the manufacture of scaffold geometries for bone tissue engineering applications. Porous scaffold geometries with open-cell structure and relative density of 10-60 v% were computationally designed and fabricated by selective laser sintering using polyamide powder. Strut and pore sizes ranging from 0.4 - 1 mm and 1.2 -2 mm are explored. The effect of process parameters on compressive properties and accuracy of scaffolds was examined and outline laser power and scan spacing were identified as significant factors.

The effect of process parameters on compressive properties and accuracy of scaffolds was examined and outline laser power and scan spacing were identified as significant factors. In general, the designed scaffold geometry was not accurately fabricated on the micron-scale. The smallest successfully fabricated strut and pore size was 0.4 mm and 1.2 mm, respectively. It was found that selective laser sintering has the potential to fabricate hard tissue engineering scaffolds. However the technology is not able to replicate exact geometries on the micron-scale but by accounting for errors resulting from the diameter of the laser and from the manufacturing induced geometrical deformations in different building directions, the exact dimensions of the manufactured scaffolds can be predicted and controlled indirectly, which corresponds favorably with its application in computer aided tissue engineering.

Introduction Tissue engineering aims to produce patient-specific biological substitutes for damaged or lost tissue. One of the principal methods in tissue engineering is to grow the relevant cells in-vitro or in-vivo into the required 3D shape. But cells can only migrate randomly in 2D layers on a surface, therefore 3D porous matrices (scaffolds) with high surface to volume ratio to which the cells can attach and differentiate have been introduced [1-2]. The scaffold aims to mimic the function of the extra cellular matrix (ECM) of the tissue to provide a temporary template for tissue growth. Cells then adhere to the scaffold, proliferate and secrete their own ECM [1-2].

Cells then adhere to the scaffold, proliferate and secrete their own ECM [1-2]. The scaffold has to have an open pore structure to enable tissue ingrowth, vascularisation, and diffusion, it has to serve as an adhesion substrate for the cells, provide temporary mechanical support, and replicate complex anatomical site geometries [3-5].Conventional technologies of scaffold fabrication like compression molding and particulate leaching, fiber bonding, gas foaming, phase separation, solvent casting and freeze drying often involve the use of porogens and toxic solvents [6]. Many of these conventional scaffold fabrication techniques can only produce scaffolds with limited thickness and random internal architecture that is not reproducible [7]. SLS and other Solid Freeform Fabrication (SFF) technologies are able to overcome these limitations and enable the design and fabrication of scaffolds with predefined internal and external architectures.

SLS is an additive particle bonding technique, where a laser beam is used to selectively bond a thin layer of powder particles together [6]. The interaction of the laser beam with the powder raises the powder bed temperature to the melting point and fuses the particles together. When one layer is sintered, a new layer of powder is deposited on top of the existing layer by a roller. This sintering process is repeated until the 3D part is ready. A key advantage of this technology is that a large variety of materials includingbiomaterials can be used given that they are available in the form of powder and are not degraded by the laser beam during sintering.SLS of tissue engineering scaffolds and other porous structures for biomedical applications has been investigated by several groups worldwide. Two main approaches are used to manufacture porous matrices for tissue engineering applications.

Two main approaches are used to manufacture porous matrices for tissue engineering applications. One uses the fact that laser sintered parts have unintentional, manufacturing-induced porosity that can be governed to some extent by process parameters, particle size and other characteristics of the powder, but the distribution and architecture of the pores are random. Using this method the fabrication of different biocompatible and bioresorbable polymer and polymer/ceramic composites into scaffolds was carried out.The second approach that will be followed in the present study fabricates scaffolds with complex predesigned internal architectures [8-12]. Williams et al. [8] manufactured scaffolds from Poly(ε- caprolactone) powder with 100μm particle size.

Williams et al. [8] manufactured scaffolds from Poly(ε- caprolactone) powder with 100μm particle size. Porosity of the designed scaffolds were not accurately reproduced, designed solid regions were incompletely dense with an average porosity of 20%, additionally excess powder was sintered and bonded to the surface of the pore architecture resulting in inaccuracies. The smallest successfully fabricated pore size was 1.75mm in diameter. Partee et al. [9] investigated five process parameters, namely laser power, scan speed, scan spacing, part bed temperature and powder layer delay time at two levels to optimize the parameter settings for the above experiment by applying a 2 level factorial design of experiments. With the found optimal settings the authors could achieve 94% density at the designed solid regions and 3-8% of dimensional accuracy; they were also able to remove most of the excess powder. Das et al. [10] selective laser sintered periodic and biomimetic scaffolds using Nylon 6.

Das et al. [10] selective laser sintered periodic and biomimetic scaffolds using Nylon 6. Periodic cell based designs with 800μm orthogonal channels, 1200µm pillars, and void fraction of 53.7 % were successfully fabricated using powder with 10-100µm particle size. The biomimetic micro-architectures with 10-100µm features relying on data from µ-CT or µ-MRI of human trabecular bone had to be scaled up (4×) to be able to fabricate them from powder with 38-42µm particle size. Tan et al. [11] fabricated 3D PEEK (25μm) / HA (60μm) composite scaffolds with strut length of 1.5 mm, pore size of 600μm, and porosity less than 80% with microporosity still present within the struts. It was found that microporosity of the struts decreased with increasing HA content.In the present study the second approach was adopted, that is, the internal architecture of the scaffold geometry is pre-designed. This paper maps the mechanical properties of scaffolds with different relative densities.

This paper maps the mechanical properties of scaffolds with different relative densities. The effect of different process parameters on accuracy and compressive properties are examined in order to identify the smallest attainable feature size. Both mechanical properties and accuracy are to be examined, since both are of significant importance to the fabrication of successful tissue engineering scaffolds. A small change in the geometry of the scaffold can influence cell interaction during cell culturing, furthermore mechanical properties are also dependent on the geometry of the cellular solid, since even at a given relative density level geometries with different features result in different compressive properties [12]. Therefore accurate geometries with precisely controlled feature shape and size should be manufactured.Materials and Methods Polyamide powder marketed under the brand name DuraForm PA was used for this study.

DuraForm PA is a white, odorless thermoplastic powder with average particle size of 58µm and rated particle size distribution in which 90% of all particles are between 25-92µm, these values were verified for post-processed powder using a Malvern Mastersizer particle size analyzer.DuraForm PA is a semi-crystalline polymer with a melting temperature of 184°C. It can be sterilized in an autoclave and is compliant with USP class VI testing. Sinterstation 2500plus with a low power, λ=10.6 µm continuous wave, CO2 laser focused to a 400 µm spot was used to fabricate the scaffolds. Prior to the design of experiments, calibration of the SLS system was performed using a millimeter – scaled calibration part and default parameter settings (Table 1a) to identify the necessary shrinkage compensations. The calibration indicated that a shrinkage compensation of 2.9 % is necessary in the X and Y direction, and 2.8 % in the Z direction.

The calibration indicated that a shrinkage compensation of 2.9 % is necessary in the X and Y direction, and 2.8 % in the Z direction. After SLS processing was completed, the scaffolds were allowed to cool overnight in the build chamber and were removed from the part bed. Excess powder was brushed off from the exterior and the internal architecture was cleaned using compressed air with a pressure of 2-6 Bars.Table 1 (a) Default parameter settings for calibration parts, (b) Factors and factor levels used in the Taguchi arrayParameter Laser fill power Scan count (fill) Outline laser power Scan count (outline) Scan spacing Layer thickness Preheat temperatureLevel 5W 1 11W 1 0.15mm 0.1 mm 143 CParameter Laser Fill Power Outline laser power Scan spacing Strut sizeLower level 6 W 0 W 0.1 mm 0.6 mmHigher Level 16 W 10 W 0.2 mm 0.8mmThe present study carries out 3 sets of experiments.

It investigates the effects of relative density (1) and process parameters (2) on the mechanical properties of the scaffolds, and the effect of process parameters on the accuracy (3) of the manufactured parts.Dependence of mechanical properties on relative density Applications of scaffolds as temporary bone substitutes in the field of tissue engineering cause them to be loaded in compression, therefore only compressive mechanical properties are studied and evaluated. Disk-shaped scaffold structures (Fig. 1a) with orthogonal channels, designed in SolidWorks 2005 and imported to STL file format were used to investigate the compressive mechanical properties of cellular solids with different relative densities. The strut and pore sizes were alternated between 0.45 – 1 and 1.4 – 1.95 mm respectively to achieve different relative density levels, the range of 40 – 10 v% was examined in 5 v% increments.

The strut and pore sizes were alternated between 0.45 – 1 and 1.4 – 1.95 mm respectively to achieve different relative density levels, the range of 40 – 10 v% was examined in 5 v% increments. In order to treat the scaffold structure as a continuum, it had to be ensured that the ratio of specimen diameter to unit cell size was less than 20 [13-14]. At each relative density level five replicates were manufactured at default parameter settings.Specimens were compressed to failure using an Instron 4202 universal electromechanical testing machine at a rate of 1 mm/s using 50 kN load cell. Compressive Young’s modulus and yield strength were determined in accordance with ISO 604:2002 standard. Since in the present experiment series, polyamide was used, which is not a potential material for tissue engineering applications due to lack of bioresorbability, normalized values are being displayed and analyzed.

Solid control specimens were manufactured in order to be able to normalize the calculated compressive properties.Dependence of mechanical properties on process parameters It is assumed that at a given relative density level the mechanical properties of the specimens can be improved by changing the parameter settings. The examined process parameters were scan spacing, that is the distance between two parallel laser scan lines, laser fill power, that is thepower in watts at which the scan lines are sintered, and outline laser power, that is the power in watts at which the contour of the object is sintered, additionally the strut size of the scaffold was also included in the design.In order to reduce the number of treatment combinations Taguchi’s L8 array was used as a one- half replication of 24 full factorial design of experiments to identify the process parameters and interactions that have a significant effect on the mechanical properties.

This array enables the examination of 4 factors at two levels including the effects of three interactions. Table 1b summarizes the factors and factor levels that were examined. Analysis of variance (ANOVA) was accomplished using Design-Expert 7.1 demo version to find the factors significantly affecting the compressive properties of the scaffolds. Disk – shaped lattice geometries (Fig. 1a) with 0.25 relative density were employed to conduct the experiment. Three replicates were fabricated from each treatment combinations, and specimens were compression tested as described above.Figure 1 Geometries designed for mechanical testing (a) and accuracy analysis (b)Dependence of accuracy on process parameters Accuracy of the manufacturing process was investigated using lattice like scaffolds (Fig. 1b). Geometries (Fig. 1b) were designed with strut sizes of 400μm, 500μm, 600µm, 700µm and 800µm in each of the tree axis planes (XY, XZ, YZ). This resulted in fewer test specimens.

1b). Geometries (Fig. 1b) were designed with strut sizes of 400μm, 500μm, 600µm, 700µm and 800µm in each of the tree axis planes (XY, XZ, YZ). This resulted in fewer test specimens. The parts were designed in SolidWorks 2005 and exported to the SLS system via the STL format.The widths of struts were measured from two directions. For example, the widths of the struts built in the Z direction were measured in both the XZ and YZ plane. A micrograph of the strut was taken in both planes using Reichert MeF2 inverted light optical microscope that was then analyzed using Buehler Omnimet Enterprise image analysis software. Grayscale thresholding was executed on the micrographs to separate the struts from the background by representing the two different ranges of grayscale values with different color bitplanes. Binary operations were applied to improve image quality and remove artifacts.

Binary operations were applied to improve image quality and remove artifacts. Once the borders of the struts were accurately mapped the mean width of the strut was calculated.Since the goal of the experiment is to identify the process parameters having significant effect on the accuracy of the struts built in X, Y and Z direction, factorial design of experiments was used to examine the effect of 3 process parameters, namely laser fill power, outline laser power andTable 2 Values of process parameters used in the factorial design of experimentFactors Strut Size Manufacturing direction Measuring direction Fill Laser Power Outline Laser Power Scan spacingLevel 1 0.4 X Side1 6 0 0.1Level 2 0.5 Y Side2 16 10 0.2Level 3 0.6 Z - - - -Level 4 0.7 - - - - -Level 5 0.8 - - - - -scan spacing. Table 2 indicates the factors and factor levels of the general factorial design of experiments setup.

Table 2 indicates the factors and factor levels of the general factorial design of experiments setup. The design has 240 (5×3×24) different treatment combinations and 3 replicates were manufactured from each treatment combination. The process parameters that are not included in the design were used at their default value.

The process parameters that are not included in the design were used at their default value. Analysis of variance (ANOVA) was carried out to identify the significant effects of the parameters on the accuracy and compressive properties of the manufactured struts.Results and DiscussionsDependence of mechanical properties on relative density Gibson and Ashby discussed [15] that three distinct regions of behavior can be identified as a cellular material is compressed and it was found that the designed scaffolds behave in the exact the same way.Figure 2 (a) Stress strain map for scaffolds manufactured by SLS, (b) relationship between the relative density and the normalized modulus (continous line – measured data, dotted line – Gibson-Ashby model), (c) relationship between the relative collapse strength and the relative density, (d) relationship between the densifications strain and the relative density of the scaffoldsThe stress-strain response of the fabricated scaffolds (Fig.

2a) is shown at different relative density levels, ranging from 0.1-0.4. First a linear elastic slope that is characterized by elastic compression of the struts can be observed at all levels, that is followed by the elastic/plastic collapse region, when the struts buckle and plastic hinges are formed in the structure, further stress compresses the solid material itself, resulting in steep densification. An increase in relativedensity of the structure improved the Young’s modulus (Fig. 2b), the increase can be described by Eqs. 1.11kssCEE(1)According to Gibson and Ashby’s model compressive mechanical properties would improve quadratically (k1=2) with decreasing porosity. However, measured data deviated from this model.

However, measured data deviated from this model. For lower relative densities, when the strut sizes were smaller measured values were significantly higher than predicted by the model, and for high relative densities, when strut sizes were greater measured values were observably lower than predicted resulting in a slower increase (k1=0.5). This deviation from the model can be attributed to the strut size effect, this will be described in more detail in a later section. As smaller struts appear more dense they therefore have significantly better mechanical properties. C1 is a constant relating to the geometry and in case of the investigated specimens it was found to be 0.3. This value accounts for both the pre-designed structure and the unintentional manufacturing induced porosity that is present in the designed solid regions.The increase in relative density elevated the collapse stress as shown in the collapse region of the stress-strain map (Fig. 2a) and in Fig. 2c.

2a) and in Fig. 2c. The relationship between the collapse stress and the relative density of the structure is given by Eqs. 2.22ksyscollapseC(2)The constants fitting the examined data are k2=2.5 and C2=0.8, again the latter constant summarizes the effects of both the designed and manufacturing induced geometry. The relative density increase also resulted in reduced densification strain, as shown in Fig. 2d.As the relative density increases the densification begins at a lower strain level. The decrease of densification strain as a function of relative density can be described by Eqs. 3 applying the constant C3=1.7.sDC31(3)As the application of scaffolds is in physiological environment, it must be pointed out that physiological strains are below 1-2%. Therefore from a tissue engineering point of view only the linear elastic region and the collapse stress are relevant, since in case of the examined scaffold structures collapse stress was reached around 2-3% of strain.

The above equations govern the behavior of the SLS fabricated scaffold with respect to relative density; however it is hypotized that by altering process parameters at a given relative density level the mechanical properties can be significantly changed.Dependence of mechanical properties on process parameters To identify the process parameters that have a significant effect on the compressive properties of the scaffolds laser fill power, outline laser power and scan spacing were simultaneously and systematically varied according to Taguchi’s L8 array. Table 3 summarizes the results of the statistical analysis after pooling the most insignificant factors as errors. The F value compares the variance of the model/parameter with the variance of the residuals (difference between experimental data and predicted values). The p value indicates the probability of the model/factorerroneously being identified as significant.

Strut s. Residual Total354.5 100.1 18.4 7.2 228.7 15.7 370.374 1 1 1 1 3 788.6 100.1 18.4 7.2 228.7 5.216.8 19.0 3.5 1.3 43.40.021 0.022 0.158 0.323 0.007Model Outline Scan s. Stut s. Residual Total40.6 25.2 7.4 7.8 2.4 43.13 1 1 1 4 713.5 25.2 7.4 7.8 0.621.9 0.006 41.0 0.003 12.1 0.025 12.7 0.023ANOVA confirmed outline laser power and strut size as influential factors for the Young’s modulus, and outline laser power, scan spacing and strut size for compressive yield strength. None of the interactions that had been included in the Taguchi array were detected to have any impact on the compressive mechanical properties.

None of the interactions that had been included in the Taguchi array were detected to have any impact on the compressive mechanical properties. The correlation coefficient, which measures the strength of a linear relationship between the experimental data and the predicted values of the regression model for the Young’s modulus and the compressive strength, was 0.84 and 0.90 respectively.Figure 3 Influence of the design parameters on the mechanical performance (a) Young’s modulus and (b) yield strengthContrary to expectations, at this size level where scan line length is within the range of the laser beam diameter, no effect that can be attributed to the laser fill power was identified. The load bearing struts of the designed disk-shaped scaffolds were fabricated in the Z direction, where the scan line length was 0.6 – 0.8 mm compared to the laser beam diameter of 0.4 mm.

The load bearing struts of the designed disk-shaped scaffolds were fabricated in the Z direction, where the scan line length was 0.6 – 0.8 mm compared to the laser beam diameter of 0.4 mm. It is assumed that the above scan line lengths (0.4 mm) did not trigger the laser.However outline laser power was detected as a strongly determinative factor, its increased value considerably improved mechanical properties of the tested scaffolds. Increased outline laser power shifts the border of linear elasticity and plastic collapse to the right on the stress-strain map (Fig.

Increased outline laser power shifts the border of linear elasticity and plastic collapse to the right on the stress-strain map (Fig. 2a) as the linear elastic region becomes steeper and the collapse stress increases resulting in improved modulus and strength values.Also, measurably improved properties were observed in smaller strut sizes and it is presumed that outline laser power is responsible for the strong strut size effect as scanning electron micrographs revealed that in the struts built in the Z direction only the outline contour was successfully fabricated (Fig. 4). When the same outline laser power was applied on the two different strut sizes (0.6 and 0.8 mm) the effective laser diameter was the same (average 180 μm at 10 W), butthe actual scanned area relative to the whole cross-section of the strut was approximately 20 % higher in the smaller struts (Fig. 4 a, b), this could explain the significant increase in compressive properties for the smaller struts.

4 a, b), this could explain the significant increase in compressive properties for the smaller struts. In struts where the fill laser scan vector length is close to the diameter of the laser beam it would seem that the laser is not triggered. However the outline scan vector is triggered and because the size of the strut is close to the diameter of the laser beam this seems to result in a densely sintered strut, for small strut sizes where the laser beam would overlap as it scans around the outside of the strut. In the case of larger struts, where the fill laser vector has not been triggered, the outline scan vector results in a hollow strut. Furthermore, as previously shown by Cheah et al. [16] when a single line is sintered the start and end portion of the line will be more dense that is caused by the extra deposited surface energy due to the acceleration and deceleration of the laser beam (dwell time).

This can also be a possible explanation for the improved mechanical properties of the scaffolds with smaller struts and therefore with shorter scan line lengths.Figure 4 Cross-section of struts fabricated in the Z building direction (a) 0.6 mm, (b) 0.8 mm strut at POl=10 WScan spacing was identified as a process parameter that significantly contributes to the yield strength of the scaffolds indicating the importance of the strength of the structure at the intersection of X, Y and Z struts, where plastic hinges are formed. If scan spacing is reduced the scanned areas will receive significantly higher levels of energy density, which results in the fabrication of denser struts in the X and Y direction, and therefore the collapse stress of the scaffold is elevated. It means that decreasing scan spacing has the same effect on the stress-strain map (Fig. 2a) as increasing the relative density of the scaffold.There were a number of unexpected results.

It means that decreasing scan spacing has the same effect on the stress-strain map (Fig. 2a) as increasing the relative density of the scaffold.There were a number of unexpected results. It was observed that the outline contour in the cross section of the struts built in the Z direction was always present, even in build parameter combinations where the outline laser power was disabled and set to 0 W. However, we were not able to verify that the outline contour is scanned when outline laser power is turned off since if both outline and laser fill power were set to zero, no parts were fabricated.In the present study only one fabrication direction was examined and only the struts built in the Z direction were loaded.

In further studies the effect of different fabrication directions should also be investigated.Dependence of accuracy on process parameters Accurate reproduction of the designed scaffolds is essential in order to ensure that the predesigned geometrical features favorable to bone cells during cell culturing and the predicted mechanical properties can be precisely reproduced. Commercial SLS machines are not designed to fabricate geometries with micron-scale features. The smallest attainable feature size is limited by the diameter of the laser that is 0.4 mm but in many cases the scan lines of the 0.4 mm struts did not trigger the laser and these struts were not formed.In general the geometry of the designed scaffolds was not properly reproduced. All the struts incorporated to the geometry were designed to be prisms with square cross-sections, but none of the fabricated struts reproduced this cross section.

All the struts incorporated to the geometry were designed to be prisms with square cross-sections, but none of the fabricated struts reproduced this cross section. Figure 5 shows the typical cross-section of the struts that was attained in the 3 building directions. The designed square cross section has been replicated as a semi-circular (half-moon) cross-section in the X and Y direction and as a circular cross-section in the Z building direction.

The designed square cross section has been replicated as a semi-circular (half-moon) cross-section in the X and Y direction and as a circular cross-section in the Z building direction. No significant difference has been detected between the dimensions of the Z struts in the ZY and ZX plane, but the dimensions of the X and Y struts in the XY plane were considerably greater than in the ZX and ZY plane, for the that reason measurements from two sides have been included in the design of experiments.Figure 5 Manufacturing induced deformations in the cross section of the struts in the different building directionsHeat conduction of the powder bed and the laser intensity change along its radius can account for the formation of the semi-circular cross sections in the X and Y direction.

Once a layer is sintered, its top surface is plane but the bottom surface curves due to the excess powder sintering around the intentionally sintered layer that distorts the designed geometry in the XZ and YZ plane. Additionally, at the examined size-scale the effect is more pronounced since the laser has a Gaussian intensity distribution along its diameter that results in a strongly curved bottom surface since the scan length is less than two-times the laser diameter. Shrinkage is also present in these struts, less observable than the curvature in the bottom surface but their top surface also becomes concave. The cross sectional difference in the struts manufactured in the Z direction can be explained by the outline scan. In the SEM images (Fig.

The cross sectional difference in the struts manufactured in the Z direction can be explained by the outline scan. In the SEM images (Fig. 4) only the outline scan can be observed, it indicates that the laser might not be triggered when the scan length is between 0.4 and 0.8 mm.The deviation from the target dimension was between -25% and +35% in the X struts and -5 and +90% in the Y struts, lower values in this range represent the measurement from the XZ and YZ plane respectively, while greater values are from the measurements taken in the XY plane. Deviation of the Z strut dimensions from the designed value was between -10 and 36% in both measuring directions.Struts of 0.4 mm in size were not fabricated in most cases when the outline laser power was turned off, if fabricated there was observable damage; therefore this strut size was not included in further statistical analysis.

By removing the 0.4 mm strut sizes from the model, residuals were significantly reduced and the remaining data better fit the regression curve.Table 4 presents the ANOVA summary for the accuracy analysis after backward elimination regression with set to 0.1. Build and measure directions are not real design factors; they were added to the model to be able to include and emphasize the uneven accuracy of the fabricated scaffolds in the different building and measuring directions and to avoid associating the parts with an average value that is less descriptive.

Strut size is not a real design factor either, but different levels of this factor were included to verify the significant parameters on the different strut size geometries.ANOVA detected outline laser power and scan spacing as significant process parameters, one- way interaction between scan spacing and laser fill power were found to be significant, therefore only these parameters and interactions were further analyzed.Table 4 ANOVA tables for accuracy (SS - sum of squares, df – degrees of freedom, MS – mean of squares, F – MSmodel/MSresidual, F value, p – probability of seeing the observed F value if there is no factor effect)SourceSSDfMSFPModel36134592613897943< 0.0001A-Strut size B-Build direction C-Measuring direction D-Laser fill power1520237 647034.9 301071 69653 2 1 1506745 323517 301071 6965159 101 94 2< 0.0001 < 0.0001 < 0.0001 0.1410E-Outline laser power F-Scan spacing AB AC BC BE BF CE435057 279874 93098 29152 58047 57897 23596 349641 1 6 3 2 2 2 1435057 279874 15516 9717 29023 28948 11798 34964136 87 4 3 9 9 3 10< 0.0001 < 0.0001 0.0001 0.0301 0.0002 0.0002 0.0266 0.0011DF Error/Residual Total126461 525312 41387721 165 191126461 318339< 0.0001Influence of scan spacing is more pronounced when the laser fill power is used at its higher level, that is 16W, it explains the significant influence of the one-way interaction between the scan spacing and the laser fill power (DF).Scan spacing has a constant size decreasing effect on all struts; meaning that as the scan spacing is increased from 0.1 to 0.2 mm, the dimensions of the struts decrease.

When accuracy is expressed as percentage deviation from the designed dimension, than scan spacing had a positive effect on the struts manufactured in the Y and Z direction; however it had adverse effect on the X struts, negative in the XZ and positive in the XY plane. The higher energy density supplied to the area in case of lower scan spacing could be an explanation for these accuracy results. However this explanation suggests that laser fill power should also have the same effect on accuracy, but this factor was found to be insignificant at p=0.14 (Table 4). The importance of outline laser power with respect to the accuracy of the specimens was expected in all directions since the outline contour has an important role in determining the dimensions of the struts.

Outline laser power had a constant size increasing effect that in terms ofdimensional deviation was negative on the struts manufactured in the X and Z directions and again adverse on the X struts in the different measuring directions. The width of these struts significantly exceeds the designed width in the XY plane and is measurably below this value in the XZ plane, which is the consequence of the manufacturing induced deformation in the cross - section (square reproduced as half moon) of these struts.Figure 6 Process parameter effects on different strut sizes in different building and measuring directions, strut size range displayed is from 0.5 – 0.8 mm, target is being 0% of deviationConclusions and Further directions This study investigated the SLS process as a potential manufacturing technology for fabricating hard tissue scaffolds. Mechanical behavior of scaffolds with open-cell internal architecture as a function of relative density was investigated.

Mechanical behavior of scaffolds with open-cell internal architecture as a function of relative density was investigated. As a result of the strut size-effect, specific for the manufacturing technology, deviations from prior described model for cellular structures were observed. The dependence of the mechanical properties on the porosity and on processing parameters, namely laser fill power, scan spacing and outline laser power was examined. Scan spacing and outline laser power were detected as significant factors in determining the compressive properties of the manufactured parts. Accuracy of the manufactured struts of 0.4 – 0.8 mm was examined in the different building directions. The cross-sections of the fabricated struts was not properly reproduced, instead of the designed square cross-section they had a semi- circular cross section in the X and Y building directions and a circular cross-section in the Z direction.

Other geometries should be examined in order to identify the features that suit the SLS technology the best. It was found that outline laser power and scan spacing are the factors responsible for the accuracy of the scaffolds. The effect of process parameters on the compressiveproperties and on the accuracy were only examined at two levels. In further experiments these parameters should be examined at more levels to enable evaluation of not only linear but higher order effects and proper optimization of the process.

In further experiments these parameters should be examined at more levels to enable evaluation of not only linear but higher order effects and proper optimization of the process. The effective diameter of the laser should be examined at different laser fill/outline power levels, since the effective laser diameter in part seems to be determining the actual size, density and therefore the mechanical properties of the fabricated scaffolds.Acknowledgements This research has been supported by a Marie Curie Early Stage Research Training Fellowship of the European Community's Sixth Framework Programme under contract number MEST-CT- 2005-020621.References 1. BL Seal, TC Otero, A Panitch, Polymeric biomaterials for tissue and organ regeneration, Materials Science and Engineering: R: Reports,31,2001,pp147-230 2.

BL Seal, TC Otero, A Panitch, Polymeric biomaterials for tissue and organ regeneration, Materials Science and Engineering: R: Reports,31,2001,pp147-230 2. MH Sheridan, LD Shea, MC Peters, DJ Mooney, Bioabsorbable polymer scaffolds for tissue engineering capable of sustained growth factor delivery, J Controlled Release,64,2000, pp91-102 3. KF Leong, CM Cheah, CK Chua, Solid freeform fabrication of three-dimensional scaffolds for engineering replacement tissues and organs, Biomaterials,24,2003,pp2363-78 4. SJ Hollister, RD Maddox, JM Taboas, Optimal design and fabrication of scaffolds to mimic tissue properties and satisfy biological constraints, Biomaterials,23, 2002, pp4095-103 5. RFS Lenza, WL Vasconcelos, JR Jones, LL Hench, Surface-modified 3D scaffolds for tissue engineering, J Mater Sci - Mater Med,13,2002,pp837-42 6. W Yeong, C Chua, K Leong, M Chandrasekaran, Rapid prototyping in tissue engineering: challenges and potential, Trends in Biotechnol,22,2004,pp643-652. 7.

CURACY EFFECTS OF SHELLING A PART IN THE SLS PROCESSN. VolpatoT ,@, T.H.C. Childs@ and A. de Pennington@T CNPq Scholar, CEFET-PR, Brazil @School of Mechanical Engineering, University of Leeds, LS2 9JT, UKAbstractIn order to reduce SLS process time in the manufacture of a mould insert, the idea of shelling the geometry of the insert has been tested. Some shelling strategies have been successful with the RapidToolTM process, proving the feasibility of the idea. It has been observed in the tests, for both polymer and RapidSteel2.0TM materials, that size accuracy, particularly of small features in the scanning (X) direction, depends on vector length (VL). When a sudden change in VL occurs, this leads to steps on the sintered surface.

When a sudden change in VL occurs, this leads to steps on the sintered surface. This paper presents both experimental observations of this and simulation results from a finite element model.IntroductionAlthough Selective Laser Sintering (SLS) is one of the technologies known as ‘rapid’, if the part to be manufactured is a mould insert, a long time can be expended in the processing. This is observed mainly because a mould insert is usually a solid part, in contrast with most plastic parts, which are like a shell. To reduce process time, it was suggested to shell the geometry of the insert, reducing mainly the volume of material to be fine/normally processed [1]. The indirect rapid tooling process called RapidToolTM, using RapidSteel2.0TM (RS2), was used to test this idea.One of the shelling strategies tested successfully required, however, using a thin shell (thickness equal to or less than 2mm) and, in this range, a side effect on accuracy of the insert was observed.

Basically, steps appeared on the walls perpendicular to the X direction, which is the main laser moving direction in a raster scanning strategy (Figure 1). In the particular case of Figure 1, a shell of 2mm presented well-defined steps up to 0.1mm inwards (undersize). In other tests using 4.5mm shell thickness, the steps were outwards (oversize). Although the size of the steps seems small, they are undesirable and will demand extra work during the surface finishing stage. If not removed, the steps can be mirrored onto the moulding or create undercuts that will make moulding ejection difficult.The steps occurred exactly where the vector length (VL - scan line length with laser on), changes from a long to a short one, and vice-versa, shown schematically in Figure 2. Shelling an insert causes a large variation of VL all over the part geometry and this variation seems to have a negative effect on accuracy.

Shelling an insert causes a large variation of VL all over the part geometry and this variation seems to have a negative effect on accuracy. In order to have a better understanding of the VL influence on part accuracy of small features in the SLS process, two approaches have been used. Firstly, two experimental parts were designed specially to check this effect on two different materials. Secondly, a 2D SLS finite element306model was used to analyse in more details the alteration in the process conditions caused by the VL changes. The main objective of this work is to present some initial findings of this study.Influence of Geometry Changes on SLS Processing ConditionsThe influence of the geometry (i.e. VL) on part properties, mainly density and strength, has been studied and discussed by other researchers [2,3,4].

VL) on part properties, mainly density and strength, has been studied and discussed by other researchers [2,3,4]. The main point observed is that geometry variation is responsible for changing thermal gradients in the part.YStepsSection of a Solid PartSection of a Shell PartXV12Short Vector Length1V2YV1Long Vector LengthLong Vector LengthScanning DirectionV2 > V1XFigure 1 - Steps on some part walls foran insert with shell of 2mmFigure 2 - VL analysis in solid and shell section schematicallyThe energy stored at the surface of the layer during the interaction period between the laser andthe powder is usually referred to as energy density or Andrew number.AN=Laser Power(P)/[Beam Speed(BS)*Scan Spacing(SS)] [Joules/mm2](1)Williams and Deckard [4] stated that this equation represents the energy stored at the surface given constant geometry of the part.

They pointed out however that the time between successive exposures (delay), which is determined by the VL and beam speed on a raster strategy, should also be considered.Delay=VL/BS[Seconds](2)During this dwell period, energy is lost due to conduction into the powder bed and radiation and convection at the surface of the powder bed. The influence of this delay period on density and part strength are associated with the difference in maximum temperature reached when varying the delay period [4]. A longer VL means more cooling and heat lost from the material being processed and, as a consequence, lower temperature attained. Higher temperature is associated with shorter VL.Other effects of geometry variation are the possibility of non-uniform shrinkage and growth. Growth is defined as sintering of powder outside the desired part boundary, caused by heat transfer from the processed material to the surrounding support powder [2].

Growth is defined as sintering of powder outside the desired part boundary, caused by heat transfer from the processed material to the surrounding support powder [2]. The shrinkage of the RS2 is307very small (an overall amount of » 0.2%) and the growth is normally considered constant and compensated together with the beam spot radius, via the offset parameter.However, in our experimental tests of the shelling idea, the geometry was the only parameter altered in the SLS processing. Therefore, it seems to be the cause of a variation in the growth in the section, resulting in development of steps on some walls. As can be seen in Figure 2, shelling causes a significant alteration of the geometry of the section to be processed.

As can be seen in Figure 2, shelling causes a significant alteration of the geometry of the section to be processed. This observation lead to a study in more detail of the accuracy of the SLS process, mainly when processing small feature in the X direction, either in pairs (typical of a shell or hollow part) or single features (typical of metal or plastic parts).Influence of VL on Part Accuracy - Experimental StudyThe experimental study was carried out using two parts specially designed to check the dimensional variation of the small features positioned in the X direction – the main direction of interest in this work (Figure 3). The part used to test single features (left) was divided into two STL files (one with the base and the features AA and the other with BB), so that, although some of these features are positioned on the same X line, they are processed individually. The X dimensions of the small features are 0.25, 0.5, 0.75, 1, 2, and so on up to 9mm.

The X dimensions of the small features are 0.25, 0.5, 0.75, 1, 2, and so on up to 9mm. The part on the right was used to study small features when processed in pairs. This design was applied for features of 1, 2, 3 and 4mm thickness, which represent the range of interest for the shelling idea [1]. The distance between the paired features was also considered, starting from 5, 10, 20 and so on up to 100mm, for 1 and 2 mm features, and the same variation for 3 and 4mm, but starting from 10mm (external distance). All the features in both parts have the same size in Y (5mm) and Z (3mm) direction. Also, all test parts have the same base design, which was used to check accuracy of larger dimensions, not reported in this work.The parts were built using a DTM Sinterstation 2000, without beam offset or scaling. The processing parameters used for RS2 were P=20Watts, SS=0.08mm, BS=1257.3mm/s and layer thickness (LT) 0.076mm.

The processing parameters used for RS2 were P=20Watts, SS=0.08mm, BS=1257.3mm/s and layer thickness (LT) 0.076mm. Also, the same experiments were carried out with a nylon-12 polymer, Duraform (DF), in order to compare the behaviour of the two materials under similar geometry conditions. The processing parameters used for DF were P=5.2Watts, SS=0.15mm, BS=1257.3mm/s and LT=0.1mm.100BABA60Figure 3 – Geometry of the test parts – single features (left) paired features (right)308Measurement ResultsAll the parts were measured using a Coordinate Measuring Machine (CMM) and the results are presented in graphs for each material. The parts were measured in the green stage (after SLS processing) for RS2. Dimensional error is usually considered as deviation of measured dimensions from the desired or nominal dimensions. We used a similar approach in this work, but we define ‘Error’ as the difference between measured dimension and VL, due to the specific objective of this study.

We used a similar approach in this work, but we define ‘Error’ as the difference between measured dimension and VL, due to the specific objective of this study. This is what error represents in the graphs that follow. The same graphs can be plotted against delay between exposures using equation (2).Figure 4 shows the results in the X direction for small features single processed according to VL for RS2. The error in the Y direction is also presented according to VL: however it is important to remember that all features have the same 5mm Y-dimension (Figure 5). Figure 6 presents the results for pairs of features and in addition, to help the analysis, the results for single features were plotted together. The large number of points for the features processed in pairs accounts for all the distances measured. The results for the paired features are also presented in Figure 7, but according to the distance that each feature of the pair is positioned apart.

The results for the paired features are also presented in Figure 7, but according to the distance that each feature of the pair is positioned apart. Figures 8, 9, 10 and 11 present the same sequence of measurement results for DF.0.850.500.800.750.700.65Error - Single Error - Single Error - Single Poly. (Error - Single) Linear (Error - Single)0.450.400.600.550.350.30Error - Single (5mm) Error - Single (5mm) Error - Single (5mm)0.500.450.25Poly. (Error - Single (5mm)) Linear (Error - Single (5mm))0.400.2001235 Vector Length (mm)46789100123456Vector Length (mm)78Figure 4 - Error of the RS2 single featuresFigure 5 - Error of the RS2 single features– X direction– Y direction0.90Error - 50.65Error - P4mm0.850.80Error - 10 Error - 200.60Error - P3mmError - P2mm0.750.700.65Error - 30 Error - 40 Error - 500.550.50Error - P1mmPoly. (Error - P4mm)Poly.

(Error - P4mm)Poly. (Error - P3mm)Poly. (Error - P2mm)0.60 0.55Error - 60 Error - 700.45Poly. (Error - P1mm)0.50 0.45Error - 80 Error - 900.400.40 0.35Error - 100 Error - Single0.350.30 0.250.300.250.2001235 Vector Length (mm)46789100.20010203040506070Dist. Between Paired Features (mm)8090Figure 6 - Error of the RS2 paired and single features – X directionFigure 7 - Error according to the distance betweenthe RS2 paired features – X direction309910010DiscussionFigures 4 and 5 show that, considering the low shrinkage of RS2, the growth does increase as the feature dimension decreases (at least for VL>1mm), or in other words, as the delay between exposures reduces. The variation is mainly observed in the X direction, and is almost negligible for the Y direction. A significant difference in growth was observed for features of the same size, but processed in pairs and singly (see Figure 6).

A significant difference in growth was observed for features of the same size, but processed in pairs and singly (see Figure 6). The single features presented much more growth than the ones processed in pairs.0.750.700.40Error - Single (5mm)0.650.350.600.300.550.500.250.450.200.400.150.350.300.250.20Error (Observed) Error (Scaled) Error (Observed) Error (Scaled) Poly. (Error (Observed)) Poly. (Error (Scaled))0.100.050.000.150123456Vector Length (mm)7891001235 Vector Length (mm)4678910Figure 8 - Error of the DF single featuresFigure 9 - Error of the DF single features– X direction– Y direction0.800.750.700.650.600.550.500.45Error - 5 Error - 10 Error - 20Error - 30 Error - 40 Error - 50 Error - 60 Error - 70 Error - 80 Error - 90 Error - 100 Error - Single Poly. (Error - 40) Poly. (Error - 30) Poly. (Error - 50) Poly. (Error - 20) Poly.

(Error - 40) Poly. (Error - 30) Poly. (Error - 50) Poly. (Error - 20) Poly. (Error - 10)0.900.850.800.750.700.650.600.55Error - P4mm Error - P3mm Error - P2mm Error - P1mm Poly. (Error - P4mm) Poly. (Error - P3mm) Poly. (Error - P2mm) Poly. (Error - P1mm)0.400.35Poly. (Error - 60) Poly. (Error - 70) Poly. (Error - 80)0.500.450.30Poly. (Error - 90) Poly. (Error - 100) Poly. (Error - 5)0.400.250.350.200.300123456789100102030405060708090100Vector Length (mm)Dist. Between Paired Features (mm)Figure 10 - Error of the DF paired and single Figure 11 - Error according to the distance betweenfeatures – X directionthe DF paired features – X directionFigure 7 shows that the growth in paired features increases when they are positioned close together. This is more clearly observed for features of 1 and 2mm.

This is more clearly observed for features of 1 and 2mm. It can be seen that for 1mm paired features, positioned 5mm apart, the error is converging to the value for 1mm single features. It seems that after around 20-30mm apart, the error is not much affected by the distance between the paired features. One thing that catches the attention is that the 2mm paired features seem to have less error than the others do, by a rather constant difference. This effect can also be seen in the overall shape of the curves for paired features (VL=2mm) in Figure 6. This observation is in line with the undersize steps observed in the strategy using 2mm-shell thickness (Figure 1), but an explanation has not yet been found.310The results for DF should be analysed taking into account that there is higher shrinkage associated with this material (» 3.4% in X direction). In Figure 8, X-errors are plotted in two ways: as observed and after a scaling to allow for shrinkage.

In Figure 8, X-errors are plotted in two ways: as observed and after a scaling to allow for shrinkage. The growth does not increase very much with small features and even presents a reduction around 1mm. It seems that the growth does not vary in the Y direction (Figure 9), except perhaps for VL<1mm. There was also not a significant variation in growth of the features processed in pairs and singly (Figure 10), and the influence of the distance between the paired features on the error values is not clear (Figure 11). It is possible however to see that the 2mm features presented, in general, the highest error of the features analysed, which is exactly the opposite of what was observed in the RS2 tests.An odd trend can be detected in the error results for the three smallest single features (0.25, 0.5 and 0.75mm), in both direction and materials.

As can be seen in Figures 4,5,8 and 9, the error increases in a quite linear way from 0.25 to 0.75mm, having a peak in 0.75mm (highest value in the experiments), and dropping rather sharply towards 1mm. It seems unlikely that this effect is due to processing changes, but it might have been caused by some systematic error of the scanner in our machine. It is worth mentioning however that a dimension of 0.75mm would correspond, because of the usual offset and shrinkage, to a nominal dimension of » 1.17mm for RS2 or » 1.3mm for DF, values that can be found in real parts.Study Using a SLS Numerical ModelA sintering model developed at the University of Leeds [5,6] was used to study the changes in the SLS processing conditions due geometry variation. The model uses the finite element method to solve heat transfer equations and calculate temperature rise caused by a transient moving heat source.

The model uses the finite element method to solve heat transfer equations and calculate temperature rise caused by a transient moving heat source. The temperature/time history that is calculated is linked to a sintering law to calculate the density changes in the bed. These density changes are used to calculate powder shrinkage in the Z direction using mass conservation. From the density profile it is possible to evaluate the size accuracy of the part. The model is developed in 2D, limiting the accuracy analysis only to the Y direction, but allowing good insight in terms of temperature and density variation in the part. Details of the model can be found in [5,6,7]. As there is not a model defined for RS2, the model for crystalline polymer was used for this study, more specifically for DF.The model was used to simulate SLS processing of single features and also pairs of features, positioned at different distance between them.

The latter was achieved by assuming a VL that induces the same delay for a single VL as for twice the feature size, plus the distance between the paired features. The fast speed set in our machine to move the mirror when the laser is off for DF is 5080mm/s. The processing parameters were the same as used in the experimental tests.As the model does not calculate the shrinkage that occurs when the material cools down from processing temperature to room temperature, the error presented should be analysed as being only due to growth.311Numerical ResultsFigures 12, 13 and 14 present respectively the maximum temperature achieved at the top surface of the layer, the average density attained and the corresponding error values for single features in the Y direction according to VL. Finally, Figure 15 presents the error according to the distance between the paired features.DiscussionThe results confirm an increase in the maximum temperature reached and density as the VL decreases.

In this small range of VL the temperature gradient does not seem very significant, causing little increase in growth (Figure 14). However, in an extended analysis, with VL up to 100mm, the variation can reach 0.17mm. The error predicted is smaller than the experimental results (Figure 9) by a value around 0.2mm.

However, in an extended analysis, with VL up to 100mm, the variation can reach 0.17mm. The error predicted is smaller than the experimental results (Figure 9) by a value around 0.2mm. The model also predicts a slight growth variation according to the distance between the paired features (Figure 15).221Tmax (C)991Average Dens.2209902199899882189872179862169859842159832149820123 Vector Length (mm)4567890123 Vector Length (mm)456789Figure 12 - Maximum temperature according Figure 13 - Average density accordingto VL for single features –Y dir.to VL for single features –Y dir.0.30Error - Single0.290.290.280.290.280.270.280.270.26Error - P4mm Error - P3mm Error - P2mm Error - P1mm0.260.2701234567890.250102030405060708090100Vector Length (mm)Dist.

Between Paired Features (mm)Figure 14 - Error for single features according Figure 15 - Error according to the distanceto VL –Y dir.between the paired features –Y dir.312Further WorkThe 2D model does not allow an accuracy study in the X direction. Therefore, a 3D model is under development using the same 2D principle. As the variation of the growth seems more evident in RS2 than in DF, a SLS model adapted to this material might help to understand this behaviour.ConclusionsAs well as the influence of VL on part densification and strength, its effect on accuracy is also relevant to further improve SLS accuracy. The experimental results showed that for small features the growth should not be considered constant and this is more critical for the X than for Y direction.The influence of the part geometry on accuracy/growth is dependent on the material being processed and also on the arrangement of features (as tested here by processing features singly and in pairs).

These influences seem much more relevant for RS2 than for DF.Although the study originated from an observation of a particular case, the shelling idea to save time in the RapidTool process, it is important to mention that geometry variation is likely to be found in most plastic or metal parts, which reinforces the importance of a better control of its effects.References[1] Volpato, N., Childs, T.H.C. and de Pennington, A. ‘Some Effects of Shelling on the SLS RapidToolTM Process’. The 8th International Conference on Rapid Prototyping, Tokyo, Japan, June 2000. pp. 395-400. [2] Nelson, C., McAlea, K. and Gray D. ‘Improvements in SLS Part Accuracy’. Solid Freeform Fabrication Proceedings, The University of Texas, Austin, Texas, 1995, pp. 159-169. [3] Deckard, C. and Miller, D. ‘Energy Delivery for Selective Laser Sintering: Issues and Directions - A Review’. 27th International SAMPE Technical Conference, October 9-12, 1995, v.27. pp. 962- 969. [4] Williams, J.D. and Deckard, C.R.

This project will investigate the present in-plane accuracy of a particular FDM machine using the benchmark “User Part” developed by the North American StereoLithography User Group (NASUG) and show the effect of optimal Shrinkage Compensation Factors (SCF) on the accuracy of the prototyped parts.The benchmark parts were built on the FDM-1650 prototyping machine and a total of 46 measurements were taken in the X and Y planes using a Brown & Sharpe Coordinate Measuring Machine (CMM). The data was then analyzed for accuracy using standard formulas and statistics, such as mean error, standard deviation, residual error, rms error, etc. The optimal SCF for the FDM-1650 machine was found to be 1.007 or 0.7%.INTRODUCTIONRapid Prototyping (RP) is the f abrication of a physical, three-dimensional part of arbitrary shape, directly from a numerical description (typically a CAD model) by a quick, highly automated, and flexible process.

RP technology has made many contributions to the manufacturing industry mainly by reducing the time to produce prototype parts and improving the ability to visualize part geometry. The physical prototype allows for earlier detection and reduction of design errors and the capability to compute mass properties of components and assemblies [1,7]. Also, it is very advantageous to present a design in client presentations, consumer evaluations, bid proposals, and regulation certification. This concept was first introduced in 1988, and since then the growth of Rapid Prototyping has been exponential every year.RP also has some challenges that must be improved upon and one of the main challenges is part accuracy. This is the main concern of countries like Japan, which are not willing to incorporate RP to there manufacturing process [1-3].

This is the main concern of countries like Japan, which are not willing to incorporate RP to there manufacturing process [1-3]. The parts produced tend to warp and/or shrink from its given dimensions forcing the user to run several trials of a part to reach its ideal dimensions or settle for a slightly inaccurate part. Another concern is the build time; it can often take several days to prototype certain parts depending on the size.77When Rapid Prototyping was first introduced, the materials used to produce the parts had low yield strength. Through the advancement in material science, the photopolymers and thermoplastics used now have much higher yield strength. The strength of RP materials is sufficient for some applications, but does not always satisfy the strength requirements for wind tunnel testing.

The strength of RP materials is sufficient for some applications, but does not always satisfy the strength requirements for wind tunnel testing. The use of stronger materials in the future such as metal-based powders will enable RP to produce parts to satisfy these requirements.The purpose of this engineering research project is to make a contribution towards the accuracy improvement of Rapid Prototyping machines. This can be accomplished by finding the most optimal shrinkage compensation factor (SCF). The Shrinkage Compensation Factor determines the amount an object should be enlarged so that it can approach the desired dimensions of the user during the shrinking process.THEORY AND ANALYSISIn this engineering project, the rapid prototyping technology used was a Fused Deposition Modeling machine (FDM –1650), built by Stratasys.

The machine uses a 3D CAD drawing that is converted to a Stereolithography file (.STL) and then it is sliced into a series of closely spaced horizontal planes [2,7,8]. The file is then sent to the machine as a build file and production of the part begins. The machine has a build area of 10x10x10 in. and produces hard solid prototypes that may be used in many applications.The FDM technology uses a spool of ABS thermoplastic filament with a diameter of .070 inches to build a part. The filament is heated to above its melting point and then extruded through a nozzle on a delivery head and onto a build platform. When one layer is finished the platform is lowered and a new layer is formed on top of it.The overall inaccuracy of the parts being built by RP technology has been one of the major challenges that need to be overcome. Errors due to shrinkage and warpage dominate the inaccuracy of the part.

Errors due to shrinkage and warpage dominate the inaccuracy of the part. The thermoplastic ABS material used in FDM machines experiences a volume change when it is heated and then extruded onto a build platform. “Each prototype made is slightly smaller than its designed dimensions”[3,5,8]. SCF is one of the variables that can be controlled by the operator to influence the overall accuracy of the part. When prototyping, the shrinkage is compensated by multiplying the dimensions by a SCF through the build software. Through research, the optimal SCF can be found and the prototype can be built within the specified build accuracy of the machine.To test the build accuracy of a prototyping machine, a user part was developed by the North American StereoLithography User Group (NASUG). It is an in-plane benchmark part, which covers the extent of the build platform and allows for the measurement of the part’s X and Y dimensions. This particular part (Fig.

It is an in-plane benchmark part, which covers the extent of the build platform and allows for the measurement of the part’s X and Y dimensions. This particular part (Fig. 1) was chosen because of: (1) its large amount of independent surfaces, (2) a variety of balanced measurements ranging from .125 to 9.5 inches, and (3) an almost equal amount of inside and outside measurements. These three characteristics made the part ideal for performing measurements because it reduced biases on certain measurements, making all of the factors in producing the part just as important [4-5].78To calculate the SCF, 46 measurements were taken in the X and Y plane [9]. These measurements were input into an Excel program and from these points, 27 outside and 21 inside dimensions were c a l c u l a t e d.

These measurements were input into an Excel program and from these points, 27 outside and 21 inside dimensions were c a l c u l a t e d . A f t e r obtaining this data, statistical equations were used to determine the overall accuracy of the part.TechniqueFigure 1: NASUG "User Part" and the CMM measurementFirst the dimensional data is converted to the differences relative to the CAD dimensions. The the errorin each ofdimensions is calculated by,Actual error = a - c(1)where a is the measured dimension and c is the CAD Dimension.Many statistical data analysis computations were derived from the 21 inside dimensions and 27 outside dimensions. The slope of the best-fit regression line of actual error vs. ideal dimension was calculated to get an idea of the accuracy of the part and to determine the new SCF.

The slope of the best-fit regression line of actual error vs. ideal dimension was calculated to get an idea of the accuracy of the part and to determine the new SCF. The slope is determined by,Slope = (42+54)*(A+B) – (C–D)*(E –F) (2)Constant where A = Σ(Inside Ideal Dimension)*(Error), B = Σ (Outside Ideal Dimension)*(Error), C = Σ(Outside Ideal Dimension), D = Σ(Inside Ideal Dimension), E =Σ(Error Outside Dimensions), and F = Σ(Error Inside Dimensions). The constant is determined by,Constant = (42+54)*(Σ(G2)+ Σ(H2)) - (Σ(H) - Σ(K))*(Σ(M) - Σ(K)) (3) where G = Inside Measured Dimension, H = Outside Measured Dimension, K =Inside Ideal Dimension, and M = Outside Ideal Dimension.Next a y-intercept value must be determined to complete the equation for the least-square regression line for which the slope was found above.

This equation is used to find the expected error in each of the dimensions and when that value is compared to the actual error, we can determine the residual error in the part.

The equation for the intercept is given by,Intercept = (Σ(A)2 + Σ(C)2)*(Σ(E) - Σ (G)) - (Σ (I) + Σ(L))*(Σ(C) - Σ (A)) (4)Constant where I = (Outside Ideal Dimension)*(Error), L = (Inside Ideal Dimension)*(Error), C = Outside Ideal Dimension, A = Inside Ideal Dimension, E = Outside Error, and G = Inside Error.79Now with the slope and intercept defined, the predicted value for the error, given an ideal dimension, can be found by the equation,y (predicted) = mx + c (5)where m = slope, x = ideal dimension, and c = intercept.This process of finding the slope, constant, and intercept is performed for both the x and y dimensions and the two results are averaged at the end.The predicted value of error is used to determine the residual square, which squares the difference between the measured dimensions and the predicted value:Residual Square = (Difference - Predicted)2 (6) With the result from equation 6, the residual error in the part can be calculated.

The residual error defines the overall difference of the measured values from the predicted error values [6]. The equation for residual error is defined by,(ResidualError=PONM 121+++)where M = Residual Square (Inside Y Dimensions), N = Residual Square (Outside Y Dimensions), O = Residual Square (Inside X Dimensions), and P = Residual Square (Outside X Dimensions).Finally the SCF can be derived. The SCF is an indication of the numerical percentage of how small or large the user part is compared to the actual CAD dimensions. The SCF equation incorporates the initial shrinkage compensation factor, which was used in the beginning of the build process.

The SCF equation incorporates the initial shrinkage compensation factor, which was used in the beginning of the build process. The equation to determine the SCF is,SCF (X) = 1 - (1 - (Slope))*(1 + SCF(X) OLD) (8) SCF (Y) = 1 - (1 - (Slope))*(1 + SCF(Y) OLD) (9)RESULTS & DATA ANALYSISAfter several trials of making the “User Part”, FDM-1.007 1 exhibits the lowest Mean, RMS, and Residual Error indicating that so far, a SCF of 1.007 gives the lowest error and the most accuracy for the user part.A scatter graph of the difference between the measured and desired dimension (error) plotted against the desired (CAD) dimension was used in the calculation of the SCF and the Total Shrinkage (TS) for the “User Part”. The Total Shrinkage is the slope of a “least squares” regression line of the Error vs. Ideal (CAD) dimension graph. It gives an indication of how far off the SCF that was used to create the part is to the SCF that will produce the most accurate part.

Ideal (CAD) dimension graph. It gives an indication of how far off the SCF that was used to create the part is to the SCF that will produce the most accurate part. A negative TS means the part was made too small and a positive TS means the part was made too big. The lowest TS value is desired for the most accurate part. FDM-1.008-P22 had the lowest Total Shrinkage value, but it was not the most accurate part as indicated by further statistical calculations performed on it.

FDM-1.008-P22 had the lowest Total Shrinkage value, but it was not the most accurate part as indicated by further statistical calculations performed on it. As Figure 2 summarizes, the greatest total shrinkage of1FDM refers to the build method, 1.007 is the SCF, and I (if used) refers to the first measurement for the part2 This is the second part made with a SCF of 1.008, due to questionable results given by the first part80e g a k n i r h S0.4000%0.3000%0.2000%0.1000%0.0000%t n e c r e P0.1000%0.2000%0.3000%0.4000%FDM- 1.005FDM- 1.006Figure 2: Total ShriFDM- 1.007FDM- 1.008-P1FDM- 1.008-P2Part IFDM- 1.009FDM- 1.010-IFDM- 1.010-II0.2966% and 0.2744% was achieved from parts FDM-1.005 and FDM-1.010-II, which indicates that the optimum SCF is not close to those SCF. Part FDM-1.007 showed the second least shrinkage of further calculations proved this was the most accurate part.

Part FDM-1.007 showed the second least shrinkage of further calculations proved this was the most accurate part. butThe Shrinkage Compensation Factor (SCF) is related to Total Shrinkage because the value for Total Shrinkage is subtracted from the initial SCF, to get a new, more accurate SCF. As i l l u s t r a t e s , F i g u r e calculations performed on the measurements revealed that a new SCF that should improve the accuracy for a part made on a FDM machine lies between 0.65% and 0.75%.

Part FDM- 1.007 was the only part that could be made in this range and it was this part that turned out to be the most accurate.1.2000%e g a k n i r h S1.0000%0.8000%30.6000%0.4000%%0.2000%0.0000%FDM- 1.009FDM- 1.005FDM- 1.006FDM- 1.007FDM- FDM- 1.008- 1.008- P1 P2 Part IFDM- 1.010-ISCF UsedNew SCFFigure 3: Used & New Shrinkage Compensation FFDM- 1.010-IIThe mean error is calculated to get an indication of the error in the part and to determine whether the dimensions were within the machine specified accuracy [6]. The graph for mean error shows a similar trend to the Total Shrinkage graph and thus gives similar information. A positive mean error Figure 4: Summary of Mean E indicates that most of the measured dimensions were greater than the CAD dimensions, and a negative mean error indicates that most of the measured dimensions were smaller than the CAD dimensions.

The FDM parts with shrinkage factors set from 1.005 to 1.010 had mean errors ranging from –0.0119 inches to 0.0110 inches, giving an even spread about zero. As Figure 4 shows, FDM-1.007, FDM-1.008-P1, & FDM-1.008-P2, produced mean errors, which satisfied the machine specified accuracy of +0.005 in. However, the mean error can be influenced by negative and positive error values, and needs to be used in conjunction with other statistical calculations to find the best part.0.0150) n i (0.01000.0050r o r r E0.0000n a e M0.00500.01000.0150FDM- 1.010-IFDM- 1.005FDM- 1.006FDM- 1.007FDM- 1.008-P1FDM- 1.008-P2FDM- 1.009FDM- 1.010-IIPart I81The standard deviation was also calculated to get an indication of the spread of errors around the mean and really understand the meaning for the value of mean error [6]. A low mean error and a low standard deviation will provide a good indication of the accuracy of a user part.

A low mean error and a low standard deviation will provide a good indication of the accuracy of a user part. The standard deviation data varied from 0.0096 inches to 0.0246 inches with the full results represented in Figure 5. Parts FDM-1.007, FDM-1.008-P1, and FDM-1.008-P2 all had mean errors very close to zero. However, when these values were matched with the standard deviation values, FDM-1.007 had the lowest spread of errors around the mean and this is another reason for the high accuracy of this part.0.0300) n i (0.02500.0200r o r r E0.01500.0100S M R0.00500.0000FDM- 1.009FDM- 1.005FDM- 1.006FDM- 1.007FDM- 1.008-P1FDM- 1.008-P2FDM- 1.010-IPart IFigure 5: Summary of Standard Devi) . n i (Calculating Residual Error for the user part allows the measurement of the errors in the part, which are not caused by shrinkage. These are random errors in the machine, material, and measurement that are much harder to measure and control.

These are random errors in the machine, material, and measurement that are much harder to measure and control. Overall, the Standard Deviation and Residual Error (Figure 6) for each user part had similar values, indicating that the error in the part due to shrinkage is very small and it is the random errors, which contribute greatly to the inaccuracy of the part. The lowest Residual Error of 0.0096in. for FDM-1.007 was the same as its Standard Deviation, indicating the error in the part due to shrinkage was extremely small and all the errors were random errors that are harder to control.0.02500.0200r o r r E0.0150l a u d i s e R0.01000.00500.0000FDM- 1.010-IFDM- 1.005FDM- 1.006FDM- 1.007FDM- 1.008-P1FDM- 1.008-P2FDM- 1.009FDM- 1.010-IIPart IFigure 6: Summary of ResidualThe most important of all the data calculated is the RMS Error (Figure 7) as it gives a better indication of the size of the error in each part.

It is a better approximation than the mean error, which can be influenced by positive and negative error values [6]. The lowest RMS Error would indicate the most accurate part. It is FDM-1.007, which gives the lowest RMS Error value indicating its greatest accuracy and verifying the previous results, which indicated that it was the most accurate part. Part FDM-1.008-P2 had lower Total Shrinkage and Mean Error values than FDM-1.007, but its Standard Deviation, Root Mean Square Error and Residual error values verified its inaccuracy.82FDM- 1.010-II0.03).

Summary of FDM-1650 ResultsConclusions and RecommendationsThe objective for this research project was to find the optimal shrinkage compensation factor for the FDM-1650 prototyping machine and produce the most accurate part. From the data that was analyzed, the best SCF for the FDM-1650 was:SCF = 1.007, or (0.7%)There are a few recommendations arrived upon for future accuracy research on the FDM- 1650. Firstly, varying temperature and build speed during the build process could affect the accuracy of the parts. These parameters might help minimize the amount of warpage that the part undergoes due to the amount of time it is exposed to heat. Also, building different percent sizes of the part can be used as a test to compare the effects of shrinkage on a small and a large part. Another recommendation is for more accurate SCF inputs, which contain 4 to 5 decimal places.

Another recommendation is for more accurate SCF inputs, which contain 4 to 5 decimal places. This would allow for accuracy research of parts with SCF’s between 1.0065 to 1.0075, the region where the results point to as having the best SCF. The machine should also allow for inputs in both the X and Y direction instead of the one shrinkage factor that can presently be input into the83build software. The addition of these inputs and added accuracy in the decimal place should produce a more accurate part that is even closer to its desired dimensions. Finally, the support material that is currently in use with this machine should be changed because it is difficult to remove. A water-soluble support material would make it easier to remove and clean the part, saving valuable time.ACKNOWLEDGEMENTThis work was funded by a National Science Foundation (NSF) grant to Loyola Marymount University for their Research Experience for Undergraduates program.

curacy of Stereolithography Parts: Mechanism and Modes of Distortion for a "Letter-H" Diagnostic PartThomas H. Pang*, Michelle D. Guertin, and Hop D. Nguyen 3D Systems Corp., Valencia, California *(Currently at Ciba-Geigy Corp., Los Angeles, California)AbstractRapid Prototyping and Manufacturing (RP&M) users need to compare the accuracy of variouscommercially available RP&M materials and processes. A good diagnostic test for both material and the fabrication process involves a 4-inch long "letter-H" diagnostic part. This diagnostic part, known as "H-4", was developed to measure the inherent dimensional characteristics of various RP&M build materials. It is also less dependent on the calibration status of particular RP&M machines, and is excellent for the purpose of generating simple but meaningful accuracy information, which can be used to further understand the mechanism and the modes of distortion in RP&M materials.

H-4 parts were prepared and built in Stereolithography Apparatus (SLA) using Ciba-Geigy epoxy based resins SL 5170 and SL 5180, and results were compared to acrylate based SL 5149. Experimental data involving the magnitude, mechanism, and the modes of distortion for these three resins are analyzed in this paper.IntroductionThe term "accuracy" associated with some of the RP&M build materials is sometimesused without sufficient supporting evidence or data, and causes misunderstanding among users hunting for materials that are essential for their applications, especially those users who are relatively new to the industry.Mechanical, electrical, and thermal properties ofRP&M parts are quite easy tocomprehend due to their commonality with other fields and applications. However, part accuracy associated with the RP&M build materials have always been quite difficult to grasp until one sees a few distorted parts.

However, part accuracy associated with the RP&M build materials have always been quite difficult to grasp until one sees a few distorted parts. Depending on the part geometry or build direction, distortion may not manifest at all. As most RP&M users have learned from experience, the distortion nature of layer-additive fabrication is not always intuitive.If a user would like to obtain a set of statistically significant dimensional accuracy data, theUser-Part I , designed by the SLA Users Group, is suitable. However, the User-Part accuracy measurment of 170 dimensions requires a CMM, and a relatively involving and sophisticated analysis of the measured dimensions. This takes a substantial amount of time, unless the procedure is automated with a computer.In Stereolithography, one can build a Christmas-Tree diagnostic2, which is an accuracytuning tool for SLA, with the resin of interest, to obtain resin shrinkage and the line width compensation values.

Nevertheless, some of the relevant distortions do not manifest in either the User-Part or the Christmas-Tree diagnostic parts.170A cantilever pare which gives the degree of curl distortion of an unsupported cantilever, isstill relevant for SL resins in some limited applications for which support are not accessible for removal. However, since most ofthe users are supporting almost all ofthe downfacing features especially with the recent availability of various automatic support software products, some of the significance of the cantilever diagnostic is quickly being lost.One can build a straight rod and measure the dimensional error compared with the CAD dimension. This diagnostic process is too simplistic. Besides, it is quite difficult to separate the inaccuracies due to the degree of shrinkage compensation applied, and the level of calibration used for the particular SLA.

Besides, it is quite difficult to separate the inaccuracies due to the degree of shrinkage compensation applied, and the level of calibration used for the particular SLA. A good diagnostic test for resin comparison must measure the inherent dimensional performance characteristic of the SL resins, and less dependent on the particular SLA machines they are tested on.In addition, to allow for a simple comparison, the diagnostic test should be relatively small, and simple to prepare, to build, to post process, and to analyze the data, as well as to interprete the results.A smaller 2-inch version ofletter "H" diagnostic was first introduced by Joe Allison in19944 and was found to be an excellent candidate for the purpose of simple but meaningful accuracy diagnostic. Now, a 4-inch version, H-4, was developed. The diagnostic parts were prepared and built in Ciba-Geigy SL resins SL 5170, SL 5180 , and SL 5149.

Now, a 4-inch version, H-4, was developed. The diagnostic parts were prepared and built in Ciba-Geigy SL resins SL 5170, SL 5180 , and SL 5149. The diagnostic design and results for the three resins are discussed below.H-4 Dimensions and DistortionsThe 4-inch-long H-4, with 0.25 inch-cross sections were prepared. The diagnostic partrepresents a "letter H". The dimensions are described below in figure 1.0.250~Front View4.0000.750 inchI~I Side View I0.2501.7501.000 inch0.750 inchFigure 1.0.250 inchThe final distorted H-4 parts were analyzed, and five critical dimensions, H-top, B-top, Waist, AnklelFoot, and Lateral, were identified. These are listed next to the object in figure 2.171H-topILateralIB-topWaist'..... _----_. __ .. _---_._-~------_._ ..... ---_.--_.-_.~ .. ....... _ _... --_ ..... _-~---_ ...--_~... ... -Ankle~---F-oo-t--1 __-+';L- t 10.100Figure 2..

_----_. __ .. _---_._-~------_._ ..... ---_.--_.-_.~ .. ....... _ _... --_ ..... _-~---_ ...--_~... ... -Ankle~---F-oo-t--1 __-+';L- t 10.100Figure 2.. ~""'-"--.In this figure, solid lines represent the CAD dimensions, or an undistorted H-4 part as it is built on the RP&M machine, and the dotted lines represent those of the distorted part, with the distortion grossly exaggerated for purposes of illustration.Modes of Distortion in H-4Three modes of distortion were found to take place in this part:1)During SLA building (In-vat Distortion)The Waist distortion happens during SLA building. As the legs are built layer-wise fromthe bottom, and the first down-facing layer of the "seat" (bottom of the lateral part) or the "bench" is drawn, the shrinkage of the curing layer pulls the legs horizontally toward each other. This shrinkage action gives rise to the formation of the "Waist" in the final, distorted part.

This shrinkage action gives rise to the formation of the "Waist" in the final, distorted part. As the subsequent layers are cured on top of the first downfacing layer, the shrinkage forces decrease relative to the strengths of the previous layers. If the distortion is large, the waist becomes a dent. When the distortion is small and gradual, the ends of the bench becomes curved or angled with as opposed to being perpendicular. Eventually, as the final layer of the 0.250 inch "bench" top is cured, the length of the bench top (B-top) becomes very close to the nominal value.However, the actual B-top dimension depends on various factors including the cureshrinkage during laser drawing, SLA calibration, shrinkage and line width compensation factors used for SLA building. As previously mentioned, a diagnostic test for comparison of materials should focus on the distortion of the resin, in the case of SL, and not on the calibration status of the particular machine.

As previously mentioned, a diagnostic test for comparison of materials should focus on the distortion of the resin, in the case of SL, and not on the calibration status of the particular machine. The former is what we're interested in, the latter is a hardware issue. Therefore, the nominal CAD dimension is not to be used as a comparison point, but the B-top dimension will be used as an internal reference dimension. Nevertheless, the dimension of the part should not be grossly different from the 4-inch length, for comparison purposes.In this test, the distortion at each of the positions designated in figure 2 will be defined bythe difference in dimension compared to the dimension at B-top. By using the dimension at B-172top as the reference dimension for this H-4 diagnostic testing, the distortion values are truly internal, and depend minimally on the SLA on which it was built.

By using the dimension at B-172top as the reference dimension for this H-4 diagnostic testing, the distortion values are truly internal, and depend minimally on the SLA on which it was built. The argument applies similarly to other RP&M machines.The Waist distortion is an extremely relevant distortion mechanism for SLA, as well asforother layer-additive RP&M methods that incorporate dimensional changes due to physical or chemical change as the materials are laid and solidified, because this type of distortion CANNOT be easily corrected by simple shrinkage or linewidth compensation operations on the build machines.The nature of the supports also strongly affects the Waist distortion. If the supportsextend continuously from one leg to another (sideways) along the long axis, filling the gap between them, the shrinkage forces of the first down-facing layer of the lateral section may not be strong enough to move the legs towards each other.

Thus, it is essential to use supports that least restrict the legs from moving toward each other for this test.In this test, "edge profile supports" have been suppressed, and only the supports that are perpendicular to the expected distorting direction were used, allowing the legs to move toward each other in the vat with least resistance. For this experiment, the support separation was 5-mm. Also, the bottom of the legs were supported with an "+ " -shaped web support.Note: The distortion at the H-top is essentially suppressed during the SLA buildingbecause the legs are tied to the platform, and the lateral bench section is also held by supports. The H-top distorts only when the lateral section bends, as discussed in the next section.

The H-top distorts only when the lateral section bends, as discussed in the next section. However, a small, non-negligible distortion occurs at the AnklelFoot positions as the legs splay inward relative to the platform in the green state.2)After Removal from the Platform (Post-support Removal Distortion)Once the "green" unpostcured H-4 part is removed from the platform, the legs are free tosplay out. There are no longer any restricting forces exerted by the supports and the platform. This frees the "seat" portion of the H-4 part, and it undergoes curl distortion in the lateral direction, as shown in figure 2. This latent lateral curl distortion, which is essentially a creep distortionS, is a dimensional manifestation of the residual internal stress of the part. As the lateral distortion takes place, the stress correspondingly splays the H-top portion (the "arms") inward, and the AnklelFoot outward.

As the lateral distortion takes place, the stress correspondingly splays the H-top portion (the "arms") inward, and the AnklelFoot outward. The H-top distortion, therefore, is a post-building distortion.At this point, the distortion of the AnklelFoot portion now has two distortioncomponents. One is a small but non-negligible in-vat distortion due to the shrinkage of the first few down-facing layers of the "seat", explained in 1), and the other is the one due to the lateral distortion, leading to the splaying of the legs outwards. The lateral distortion generally increases over time, especially when the distortions correspondingly increase.part is left in the green state, and the H-top and AnklelFoot173For purposes of standardizing this H-4 diagnostic test, all of the dimensional measurements for the green part were taken one hour after the part was removed from the platform.

Because there is some rounding of the downfacing edges that increases measurement error at the Foot position, dimensions at the Ankle position, which is offset from the Foot position upward by 100 mils, was used for calculating the degree of splaying out. Hence, the dimensional measurements taken at the Foot position were not utilized for analysis.3)After UV Postcure (Postcure distortion)The green H-4 part is then UV postcured in an appropriate UV oven for an appropriateduration from one side only, laying down. In this test, H-4 parts were UV post cured for one hour. After UV postcure, all dimensions are expected to shrink equally. There are, of course, additional internal stresses introduced by the postcuring process, which may further distort the part. However, the expected postcure distortion is mostly in the plane of the H-4 part when postcured in this manner. The distortions in the plane of the H-4 part is not considered for this diagnostic.

However, the expected postcure distortion is mostly in the plane of the H-4 part when postcured in this manner. The distortions in the plane of the H-4 part is not considered for this diagnostic. The inhomogeneous part of the postcure distortion relative to the dimension at B-top, is expected to be small enough for this geometry such that it can be neglected for this test, as hown in the data presented. The overall postcure shrinkage, however, is included for this test. The postcure shrinkage value is also measured at the B-top position.Experimental H-4 Distortion of SL 5170, SL 5180, and SL 5149The H-4 diagnostic part was built in epoxy based Cibatool resins SL 5170, SL 5180, and acrylate based SL 5149. The HeCd and Argon ion laser powers were 20 mW, and all parts were built in 6-millayers (1.75 inch height = 292 layers). The resin specific build parameters recommended by 3D Systems were used, including shrinkage compensation and line width compensation factors.

The resin specific build parameters recommended by 3D Systems were used, including shrinkage compensation and line width compensation factors. SL 5170 and SL 5180 were built in the ACES ™ style whereas SL 5149 was built in the recommended STAR-WEAVE ™ style.All parts were built without sweeping for this experiment. Z-wait was increased to 45 seconds for the epoxy based resins, and 60 seconds for the acrylate to ensure appropriate layer recoating. Sweeping action is not expected to affect the results. However, too fast of a sweep may sway the parts back and forth in the vat if the part is built horizontally against the sweeper blade, and may result in layer-to-Iayer dislocation. Besides, sweeping introduces error associated with the blade gap adjustment inaccuracy, if it is present.Without sweeping, building three H-4 parts takes approximately 8 hours at the laserpower of 20 mW. If a sweeper is used, it should take about 6 hours to build one H-4 part.

The measurement error associated with this part, including human error, is about ±O.5-1.0 mils. The values are listed in Table A for a green postcured one. The distortion at each positions, H-top, Lateral, Waist, Ankle, and Foot, were calculated relative to the reference dimension at B-top. They are given in Table C for the green part, and in Table D, for the postcured part. Negative numbers indicate that the dimensions are less than the corresponding B-top dimension.part, and in Table B, for aH-4 Distortion Data AnalysisFirst, notice that the relative differences between the distortion values of the green and postcured parts given in Tables C and D are very small, indicating that the postcuring process leads to a relatively uniform shrinkage. The lateral distortion, however, should be considered separately because the lateral distortion changes with time, and there are inadvertently some time lapse before the H-4 parts were postcured.

The lateral distortion, however, should be considered separately because the lateral distortion changes with time, and there are inadvertently some time lapse before the H-4 parts were postcured. In this experiment, the parts were left in the green state after removal from the platform for one hour. Then they were UV postcured for one hour.Linear Green Build ShrinkageThe difference between the Waist dimension and the B-top dimension represents theshrinkage that occurs in the green state during the building process as discussed in section 1). The Waist distortion results mainly due to the shrinkage of the first downfacing skin layer of the bench section, as well as some contribution from the shrinkage resulting from a few subsequent layers. This may be referred to the "linear green build shrinkage" for a 4-inch long part. The values are shown in Table E.As accurate as the epoxy resins may be, their linear green build shrinkage is still in therange of 8-12 mils/4inch.

The values are shown in Table E.As accurate as the epoxy resins may be, their linear green build shrinkage is still in therange of 8-12 mils/4inch. This is still a few factors away from the User-Part RMS error of about . Furthermore, the User-Parts do not explicitly demonstrate dimensional inaccuracies of this type. The build shrinkages must be decreased even more, to futher increase SL part accuracy. The epoxy resins, nonetheless, have about five times less build shrinkage than the acrylate based 5149 resin.mils6Linear Postcure ShrinkageThe postcure shrinkage can be calculated from comparing the B-top dimension before andafter postcuring. This is straight forward. The values are given in Table F.Postcuring can introduce additional internal stresses. For thin sections, a nearlyhomogeneous UV postcure can be achieved.

This is straight forward. The values are given in Table F.Postcuring can introduce additional internal stresses. For thin sections, a nearlyhomogeneous UV postcure can be achieved. However, postcuring thick sections inevitably results in an inhomogeneously cured regions, due to the light absorbance characteristics of the photopolymer material. The surface will be cured more than the interior. If the material does not shrink by postcuring, no additional internal stresses will arise. Indeed, a perfect SL resin will have no postcure shrinkage.176SL 5170 and 5180 have a measurable postcure shrinkage of 0.01%. The acrylate basedSL 5149 has almost 25 times more UV postcure shrinkage, of about 0.24%, or about 10 mils/4inches. Ankle Distortion (dAnkle)As explained previously, the legs of the H-4 part splay outward due to two distortion mechanisms.

Ankle Distortion (dAnkle)As explained previously, the legs of the H-4 part splay outward due to two distortion mechanisms. One is the in-vat shrinkage distortion that snaps the two legs toward each other, and the other is the lateral curl of the bench· section that happens after removing the part from the platform. The dimensional change at the Ankle is difficult to interprete because it is a combination of two different types of distortion mechanisms, distorting in the opposite directions.The linear shrinkage, especially for the first downfacing layer of the bench section, pulls the top of the legs inward, leading to a small but non-negligible negative (narrowing) distortion between the two Ankles.

When the supports are removed, the lateral distortion generates the outward splaying of the legs, which results in a positive (widening) distortion at the Ankle position.The effective splaying out of the ankle, dAnkle eff, may be compensated by the difference between the inward distortion of the Waist, by offsetting the inward distortion distance. In other words, dAnkle effdAnkle - dWaist. The effective ankle splay values are tabulated in Table G.The legs of the H-4 part built in the epoxy based SL 5170 and 5180 splay out about anorder of magnitude less than the acrylate based SL 5149, about ~10 mils versus-IOO mils.H-Top "arms" Distortion and the Lateral DistortionThe H-top distortions (dH-top) for the epoxy versus the acrylate based SL 5149, shown in Table D, are dramatically different. The dH-top of the epoxy based resins are at least an order of magnitude less than that of the acrylate based SL 5149.

The dH-top of the epoxy based resins are at least an order of magnitude less than that of the acrylate based SL 5149. The dH-top of the epoxy resins SL 5170 and 5180 are abouf-3 mils, compared to about -30 mils for 5149.As discussed previously, the dH-top does not take place in the vat during SLA building. Itoccurs only after the parts are removed from the platform, because of the lateral distortion that happens at the bench seat section. In other words, the dH-top is also a measure of the lateral distortion (bending of the middle section). As an evidence, the dH-top values and d-Lat values scale similarly for all three resins tested. The lateral distortion phenomena in the green state is the same distortion that occurs in the green creep distortion observed elsewhere7 values for the epoxy resins indicate that the lateral bench section is dimensionally stable throughout green to postcured states..

The low dH-topThe lateral distortion, (d-Lat), may also be calculated from the measured dH-top values bythe following equation, derived trigonometrically:d-Lat ~h)dB (top X (dB-top - dR-top)( 1- cos [(1/2) (dB-top - dR-top)h])Equation 1.177where h (=0.750 inch) is the height of the "arms" above the lateral bench seat section. In this experiment, the lateral distortion was experimentally measured by the Micro-Vu. This equation assumes that the angular distortion ofthe arms are small (arcsin (alb) ~ (alb) for (alb) « 1 where a = (1I2)(dB-top-dH-top) and b=0.750 inch).For example, in the case ofSL 5170, a = -0.0012 inch, which gives (alb) = 0.00160, andarcsin(0.00160)inradians=0.00160. The comparison between the calculated and measured d-Lat values are tabulated in Table i. The agreement is very good, except for the value of the postcured SL 5149.

The comparison between the calculated and measured d-Lat values are tabulated in Table i. The agreement is very good, except for the value of the postcured SL 5149. This equation allows even those with only precision calipers to calculate the lateral distortion based on the measured dH-top dimensions. Lateral distortions cannot normally be measured using calipers.Waist DistortionThe distortion that gives rise to the formation of the "waist" on the H-4 diagnostic part, (dWaist), is perhaps the most significant of all. First, the mechanism by which the waist forms is unique to layer-additive fabrication. Second, the dWaist is the most relevant distortion for SL technology because this distortion cannot be fixed by varying the two major accuracy maximizing tools in SLA: the shrinkage compensation and line width compensationfactors.

This distortion can not easily be fixed by controlling postprocessing either, because dWaist is an IN-VAT distortion that happens in the SLA during the building process. When two legs splay inward, resulting in the wrong positioning compared to the CAD dimensions, there is not much one can do to fix it.Of course, for this particular geometry, adding more supports can restrict the dWaist.However, there are many geometries that will not tolerate such over-supporting. Most of all, one can almost be certain that resins that show high dWaist values are generate parts that are dimensionally relatively inaccurate.

Most of all, one can almost be certain that resins that show high dWaist values are generate parts that are dimensionally relatively inaccurate. The H-4 analysis can be summarized in the table below:Linear Green Shrinka e Linear Postcured Shrinka e Effective Ankle S la Lateral Curl Distortion, Green Lateral Curl Distortion, Postcured0.29% 0.01 % 14.4 mils 1.8 mils 1.4 mils0.21% 0.01% 7.3 mils 1.8 mils 2.1 mils1.16% 0.24% ":19.4 mils 17.6 mils 24.7 milsSummaryA simple diagnostic part, a 4-inch long letter-H shaped "H-4", is proposed as a tool toallow the RP&M users to compare various materials commercially available. This diagnostic part, despite the simplicity, exposes many relevant accuracy issues of the build material and the process to the users, both visually, and quantitatively.

This diagnostic part, despite the simplicity, exposes many relevant accuracy issues of the build material and the process to the users, both visually, and quantitatively. There are numerous advantages for this diagnostic part.178because it uses its own internal dimension as a reference, does not depend on thecalibration of the hardware, or the shrink factor, or line width compensation factors used for the building of the diagnostic. The distortion values can be compared with each other even if the H-4are built on different machines.The in-vat waist "tucking" distortion, latent lateral curl distortion that occurs followingsupport removal, and the splaying out of the arms and the legs are easily measured usingstandard calipers. A CMM or high precision specialty equipment are not necessary for this diagnostic, unless a more precise work is required.green build shrinkage and postcure shrinkage can be easily calculated from thediagnostic part.

Lateral curl on the horizontal section, which was not previously accessible for measurement with a caliper, can now be accurately calculated from the distance between the arms measured using calipers.diagnostic contains elements of the Christmas-Tree (linear shrinkage) and theCreepBar (latent curl distortion), and an additional in-vat diagnostic capability (waist distortion), plus visualization of distortion. The distortion appears separately as the waist distortion, and the out-vat distortion appears in the distortion of the vertical arms. The two distortions can be clearly separated in this diagostic.diagnostic has an advantage that the in-vat buildBest of all, most of the distortion can be visually inspected. If the distortion is extremelysmall, one can feel the edges of the part to determine if the waist distortion is present.

If the distortion is extremelysmall, one can feel the edges of the part to determine if the waist distortion is present. The splaying of the the arms and legs is visually detectable in most cases, because even the small lateral distortion is amplified in this 4-inch long part. If the H-4 part does not appear distorted, one can be quite sure that the particular build material will build relatively accurately.this report, thediagnostic test was applied to the Ciba-Geigy SL 5170,5180,built in ACES™style, and SL 5149 resins, in STAR-WEAVE™. The diagnostic clearly demonstrated the accuracy advantages of the epoxy based over acrylate based resins. The epoxy based only 1/5, and 1/24 compared to that of acrylate based SL 5149, respectively. Effective ankle splay distortion values, and lateral curl distortion both in the green and postcured states, are also about an order ofmagnitude less for the epoxy based resins.

These results correlate very well with numerous other diagnostic parts built at 3D Systems, as well as with the general experience5170 and SL 5180 were found to have linear green, and postcure shrinkages that areresin users, fed back from the SLA users.An accuracy study of other RP&M materials and techniques using this diagnostic will be"",vi·r""'~'Y\"",II"O interesting, and highly desired, to increase the understanding of the distortion mechanisms both during and after part fabrication.179References1 Gargiulo, Ed, Proceedings of the 1991 North American Stereolithography Users GroupMeeting, Orlando, Florida, March, 1991.2 Jacobs, Paul, Rapid Prototyping andManufacturing: Fundamentals ofStereolithography, published by the Society of Manufacturing Engineers, Dearborn, Michigan, July, 1992.3 ibid 4 Allison, Joe, Proceedings of the 1994 North American Stereolithography Users Group Meeting,Phoenix, Arizona, March, 1994.5 Pang, Thomas, "Stereolithography Epoxy Resin Development: Accuracy and DimensionalStability", Proceedings of the Solid Freeform Fabrication Symposium, University of Texas at Austin, Austin, Texas, August, 1993.6 Pang, Thomas, "Stereolithography Epoxy Resin Development: Accuracy, Dimensional Stability, and Mechanical Properties", Proceedings of the 1994 North American Users Group Meeting, Phoenix, Arizona, March, 1994.7 Pang, Thomas, "Stereolithography Epoxy Resin Development: Accuracy and DimensionalStability", Proceedings of the 1993 North American Users Group Meeting, Atlanta, Georiga, March, 1993.180.

curate Heart Model for Pacemaker Development via SFFR.S. Crockett1, T. Horvath1, M. Koch1, M. Yang21California Polytechnic State University, San Luis Obispo 2St. Jude Medical, Sylmar, CAABSTRACTMedical imaging combined with SFF techniques were used to create detailed CAD and physical heart models for commercial development of Pacemakers. Using a data set of 2D optical slice images of the human heart at 1mm spacing obtained from the Visible Human Project, a 3D CAD model was constructed by masking the features of interest in each slice. Normals on the resulting .stl file were inverted to create a single-piece mold, which was built in starch using 3D Printing. Flexible silicone was cast into this mold, and the starch was dissolved away to produce the final physical heart model.

Flexible silicone was cast into this mold, and the starch was dissolved away to produce the final physical heart model. The resulting model simulates the mechanical properties of an actual heart, with medically accurate internal and external details including major veins & arteries, coronary sinus, etc.INTRODUCTIONSolid Freeform Fabrication holds enormous promise for the field of biomedical engineering. The manufacture of objects from multiple materials with extremely complex geometries lends itself to the creation of customized implants as well as complex scaffolds, which will be key to future tissue engineering applications. In the shorter term, biomedical engineering companies are already going beyond the use of SFF for design visualization and development of surgical tools. There are numerous cases of SFF being used to create physical models of a specific patient’s anatomy. Surgeons can physically handle these models and use them to prepare for complex operations.

There are numerous cases of SFF being used to create physical models of a specific patient’s anatomy. Surgeons can physically handle these models and use them to prepare for complex operations. Also, these models can be used as patterns to make specialized implants, specific for the patient at hand.The majority of this work has involved using CT images as the input data for hard tissue such as bone. In these cases, the data segmentation is relatively simple and is generally performed automatically. When one introduces soft tissue that has less grey-scale differentiation from the surrounding tissue, the process becomes much more complicated. This paper presents a case study to illustrate both the potential and the current issues with using SFF to create geometrically and mechanically realistic physical models from medical images of soft tissue.

The human heart provides a good example, as it is both a critical organ for biomedical engineering research/product development as well as an extremely complex, organic geometry which pushes the limits of current manufacturing technology. Cal Poly has recently completed a project for St. Jude Medical to create a medically accurate rubber model of the human heart to be used for the development of pacemakers.765HEART MODEL CASE STUDYMotivationAn Implantable Cardioverter Defibrillator (ICD) is an implantable device similar to a pacemaker that is also capable of providing a defibrillating shock as necessary. The thin wires extending from the device (Figure 1), through the veins, and anchored into the heart wall are the leads. Leads are extremely complex devices, and are the focus for much of ICD product development. Product development of an ICD, like any other product, requires fit checks, mockups and prototypes.

Leads are extremely complex devices, and are the focus for much of ICD product development. Product development of an ICD, like any other product, requires fit checks, mockups and prototypes. Unfortunately, for a heart this means using the prototype device with either animal models (such as dogs) or cadavers. Both require complex procedures to use and are ultimately expensive, and neither is entirely suitable for the work associated with early product development efforts. As a result, the early design optimization of an ICD is often limited. Designs are quite mature by the time they are tested within the context of an actual heart geometry, which is clearly not the best way to do product development.What is required is a CAD model of the heart for studies such as geometry/fit checks and Finite Element Analysis of performance, and also physical models which accurately simulate the geometry and mechanical properties of the heart.

Physical models are especially desirable, as they can ultimately be used to evaluate new or novel implantation procedures for optimal ease of lead insertion. This physical model could also potentially be used in both marketing and training efforts, to help surgeons understand and evaluate the use of a new lead design or implantation procedure. For these benefits to be realized, it is important that the heart be anatomically and mechanically correct and have a realistic texture, especially the interior surface of the chambers of the heart.Figure 1. Implantable Cardioverter Defibrillator (ICD)766Data Set: The Visible Human ProjectThe data set used for the heart model was obtained from the Visible Human Project (http://www.nlm.nih.gov/research/visible), created by the National Library of Medicine (NLM).

This data set consists of transverse CT, MRI, and optical cryosection images (optical images of the sliced cadaver) of a representative male and female cadaver at one-millimeter (male) and 1/3 millimeter (female) intervals. The corresponding sections of the three imaging modalities are in the same planes, enabling comparison of individual features.For the male, the CT data consists of axial data of the entire body taken at 1 mm intervals, with a resolution of 512 by 512 pixels, each pixel with 12 bits of grey tone. The axial optical cryosection images are 2048 pixels by 1216 pixels, with each pixel containing 24 bits of color. Each cryosection image is approximately 7.5 megabytes. The cryosection cross-sections are also at 1 mm intervals, coinciding with the CT scans. There are 1871 cross-sections for both the CT scans and the anatomical images. Pixel resolution in the cross section (x-y plane) is 0.33 mm by 0.33 mm.

There are 1871 cross-sections for both the CT scans and the anatomical images. Pixel resolution in the cross section (x-y plane) is 0.33 mm by 0.33 mm. The entire human body is thus represented by a data set comprised of 5,189 anatomical images which can be viewed as 0.33 x 0.33 x 1 mm resolution voxels in 3 medical image modalities.Figure 2 illustrates the three imaging modalities. Note that MRI and Optical may at first glance look similar if both are viewed in grayscale, but the actual imaging is very different: MRI is density map, while the optical image is reflection of light. They thus pick up different features and are complementary. As can be seen in Figure 2, the heart is not very detailed in the CT images. CT is an ideal imaging technique for bones, but is poor at picking up the details of soft tissue.

As can be seen in Figure 2, the heart is not very detailed in the CT images. CT is an ideal imaging technique for bones, but is poor at picking up the details of soft tissue. MRI images would work as a data set for the heart, but the resolution of the MRI images contained in the Visible Human Project is 4mm voxels, which is too low for this effort. The optical images the heart region were thus selected as the input dataset.Figure 2. Visible Human male’s thorax. CT (top left), MRI (right), Optical Cryosection image (lower left)767Data SegmentationSegmentation is the process of differentiating the tissue of interest from the surrounding tissue, identifying and marking features of interest in each layer such that software can later interpolate between layers to create a 3D image or CAD data file.

This involves viewing each of the cross-sectional images and isolating the areas of density that indicate the type of tissue to be modeled.There are a number of commercial software packages available for both performing data segmentation and converting from a series of CT/MRI images to .stl files. MIMICS from Materialise (www.materialise.com) was used for this project. Most segmentation processes involve a thresholding technique that allows the user to highlight pixels depending on their gray scale values. For most applications involving isolation of bone, the thresholding tool is the primary tool used. Thresholding enables a mask to be created by defining a grey scale range. By specifying the upper and lower threshold boundaries, all the pixels with a grey scale value that fall within the boundary values are highlighted.

By specifying the upper and lower threshold boundaries, all the pixels with a grey scale value that fall within the boundary values are highlighted. The thresholding tool is especially useful in segmenting tissues with a comparatively low or high threshold, like bone, which show up as white in CT images.The use of optical data requires manual marking of tissue of interest, because there is no “single color” of the heart -- the heart tissue to be isolated does not differ greatly in color or brightness from the surrounding tissues, making the segmentation process more complicated. A significant amount of operator interaction was required to identify and mark the heart tissue within each optical image, layer-by-layer (~180 layers total). The pericardium (the membrane sac which contains the heart) provided a convenient visible boundary between similar looking tissue and fat, and was used as the outer boundary to define the heart.

The pericardium (the membrane sac which contains the heart) provided a convenient visible boundary between similar looking tissue and fat, and was used as the outer boundary to define the heart. The pericardium was seen as an extremely thin dark band separating layers of identically colored fat. Figure 3 shows this process, as well as the resulting .stl heart model.Figure 3. Left: Axial (left), transverse (upper right), and longitudinal (lower right) views as seen in Mimics. Green areas are areas highlighted to create a mask representing the heart model. Right: three-dimensional representation (STL file) of heart model in Mimics.768Figure 4. Coronary Sinus, seen in lower center of this view of the heart model.Note that surface of heart looks smooth, as opposed to having the expected network of veins and arteries on the surface. This is because you are actually looking at the outside of the pericardium, which is smooth.

This is because you are actually looking at the outside of the pericardium, which is smooth. The veins are present, but below the surface of the CAD model.There were some difficulties in the segmentation process. Veins and arteries were difficult to segment, because there is no clear outer boundary for the vessels, only an inner boundary. It was very challenging to obtain a consistent width of the vessel wall, and therefore to have a smooth outer boundary representing the external wall of the vessels. Also, because the initial thresholding is performed in the axial view corresponding to the orientation of the data set, the vessels that run in the other two directions were not represented as clearly as those that ran axially. The coronary sinus (lower center in Figure 4) is of great interest to developers of ICDs, but is not visible in the CAD model – it is actually just below the surface as described previously, although it is a complete, open channel.

A portion of the surface of the heart had to actually be “erased” layer-by-layer to remove the pericardium and expose the surface of the coronary sinus.After the segmentation was completed, the resulting series of layers could be interpolated in MIMICS and exported as an .stl file. A solid model was also desired to assist in ICD product development efforts (to perform simple geometry and clearance checks), and to be used in finite element analysis (FEA) studies. For these uses, the model was output as set of contours in IGES format (Figure 5), and manipulated with ProEngineer.The resulting .stl file was 80 MB, which is a difficult file size to manage. In order to reduce the file size, the resolution in the XY plane was reduced by three times. This was done by averaging over 9 pixels (0.33 x 0.33 mm) to create a representative grey scale value for the new larger pixel (1 x 1 mm). Instead of 0.33 x 0.33 x 1 mm voxels, the lower resolution file had 1 x 1 x 1 mm voxels.

Instead of 0.33 x 0.33 x 1 mm voxels, the lower resolution file had 1 x 1 x 1 mm voxels. This yielded a file that was only 14 MB, which was a more appropriate size to manipulate. The lower resolution also provided more smoothing for better on-screen visual representation. The full resolution .stl file was used for physical model creation via Solid Freeform Fabrication769Figure 5. IGES file used for finite element analysis.Physical ModelIt was determined that no existing commercial RP machine would be suitable to create the heart directly, because of the specific required elastomeric properties of the material combined with the complex geometry of the heart. Thus the selected method was to use 3D Printing to create a pattern, perform a secondary polyurethane casting process to create a mold, and then use this mold to cast the final model in a suitable RTV silicone rubber. The system used for this project was the Z Corporation ZPrinter 310 model (http://www.zcorp.com).

The system used for this project was the Z Corporation ZPrinter 310 model (http://www.zcorp.com). 3D Printing is an ideal system for a project such as this, because the unbound powder provides support for the extremely complex organic geometry.The heart has many inner chambers and an interior texture that make it impossible to mold as a single unit. Therefore, the computer model had to be sectioned into pieces that could individually be successfully molded. Even so, each of the sections still had significant overhangs/undercuts such that rigid patterns would be impossible to remove from the cast polyurethane mold material. This challenge was overcome by the use of sacrificial patterns that could be dissolved after casting the mold. Starch-based powder (ZP14), without any post- processing, was used as the build material, with the intention of dissolving it away after casting a flexible polyurethane mold around the pattern.

Starch-based powder (ZP14), without any post- processing, was used as the build material, with the intention of dissolving it away after casting a flexible polyurethane mold around the pattern. Once the mold was created and the pattern removed by dissolving, the final cast rubber heart sections were removable from the770Figure 6. Left: Exploded view of four sections. Right: Expendable patternpolyurethane mold, because both components were flexible enough that the final cast piece could be pried out of the mold. Figure 6 illustrates the chosen sections, as well as a representative starch pattern.After casting, the four resulting cast RTV rubber sections were glued together to create the final heart model. Interior and exterior detail, as well as basic mechanical properties, were sufficient for the intended use of the physical model as a tool for ICD product development. This method, however, was not entirely satisfying because of the need to create individual sections.

This method, however, was not entirely satisfying because of the need to create individual sections. As an alternative approach to casting such a complex geometry, an expendable starch mold was built directly on the 3D Printer. To create the mold geometry, the .stl normals of the heart model were flipped, and the result was Boolean joined with a block to produce the mold (Figure 7). This mold was built on the 3D Printer in the starch material, and again did not receive any post-processing so that it would be easily removable after casting the RTV.It was extremely difficult to remove the unbound powder from deep, thin cavities such as the heart walls without damaging the fragile starch mold. The results of this method of casting were promising, but showed numerous areas of missing material on the surface of the resulting RTV heart, due to unbound powder remaining in the deep, thin cavities of the mold which prevented complete filling.

Current efforts are focused on using a power/binder system which is more compatible with the methodology of directly creating expendable molds (e.g. sugar bound with a water-resistant binder).An additional ongoing study which resulted from this project is the creation of separate models of the coronary sinus along with the associated vasculature (Figure 8). Detailed physical models of vasculature systems will be useful for fluid flow studies. For example, it is possible to create “disease states” in CAD (e.g. draw a 2mm layer of plaque), produce a physical model with appropriate mechanical properties using the above method, and instrument it for use in fluid flow testing. The results of such testing would be complementary to FEA fluid dynamic analyses, and would provide access to an unlimited range of geometries to simulate age, disease, etc.Figure 7. Expendable mold.771Figure 8.

Expendable mold.771Figure 8. Isolated heart vein structure for physical testing.ACKNOWLEDGEMENTSThe authors would like to acknowledge St. Jude Medical for their support of this project.772.

curate Rapid Prototyping by the Solid Ground Curing Technologync.The first stage of Rapid Prototyping life cycle as a new technology in the marketplace is gradually ending, and the second stage has already started. Many new vendors have introduced their products in this field, utilized different, new technologies or improvements of the existing ones. Stereolithography created a stunning impression in the marketplace. After a couple of years, as customers and users have gained much experience and understanding or RP technology, the first enthusiasm started making way to more serious and demanding approach. This is very well reflected in the thorough evaluations of the different technologies available today in the marketplace, done by customers looking for a technology that will best fit their needs.

This is very well reflected in the thorough evaluations of the different technologies available today in the marketplace, done by customers looking for a technology that will best fit their needs. This is actually why most of us are here today.The first introduction of the RP concept andOne of the most commonprol:>lems;s~aredl:>YmanyRPtechnologies is model shrinkageand·dis;tortion,.cons;iderablylimitingtheaccuracy that can be achieved. .In. this;cas;ef, "aoouracyU should have onl:y one meaning: ·.the deviation of •• the·modeIJs··qimens;ionsfrom ••• the .designer' s instructions. Although there are many. interesting maohine specifi.cations that Can l:>eintroCiuceCi,liketherrepeat.al:>il.ity.of the laser beam, or the acourae;,yof.a.oe:rtain meoh.anism.inthell\achi.ne, the really important thing is the aocuraoy of. the moCiel itself. several technologies theresponsibilityt;.o the· model IS. accl.lracy is waived.

the moCiel itself. several technologies theresponsibilityt;.o the· model IS. accl.lracy is waived. to the o.perator · due to .shrinkage<effect, dis;tortionsand the need fOr supporting structure. experienced the operator is,This; actually means that the morethe better parts he will be making.Asa seconCigeneration tecbnology,the Solid Ground Curing process developed by Cubital offers s;olutions to the inherent problems of indepenqent of pres;entlyknown RP technologies, operator's s;kills. MoCiels b.uilt on the. Solider 5600 system are accurate to 0.1% of any dimension. There is no final curing, is no limit to thegeometricalshat:>esthat can be produceCi, no Sl.lpport structures are needed anCi.·several resins •. can be l.lsed by the system. for the proquctionof either one largemgCiel or many different smaller models that will be nested together anCimade in a single run. Higher safety and throughput are other benefits offered by this technology.and doitThe.

Higher safety and throughput are other benefits offered by this technology.and doitThe. spacious proCiuctionvolumeof 20" X14"X. 20" allows110Inthethereare two major principles to Cubital's Solid Ground Curingtechnology:1..The resin is exposed to UV flood light through a computerized programmable mask2..The model a vat full of liquidgenerated within a solid environment and not inThe process starts with the creation of three dimensional composition of the . t o be made in th$.next run and the definition of layer thickness. together into layers and generates a precise raster. image of each layer.The Solider computer .thenslic$s the whole compositionThis image is sent ot the mask plotting unit in the machine and a high resolution, precise optical mask is generated by means of electrostatic charges and black toner powder ..The which and undermask is then positioned precisely above the workpiec~ already been spread with a thin layer of.

liquidphotopolymer high pOwer UV lamp (200.0W) ,·readyito be exposed.opened for about 3 seconds and the resin is exposed and light passing through the mask.. of the. mOdel ·at this layer is. fully cured ·due to .. the and the length of exposure time. Unexpo.sedareasThe pattern of theremainih liquid state.mask, which is actually a plate of glass, moves back to the mask ready for the nextunit and is. being discharged and erased,to be drawn.workpiece now passes under an aerodynamic wiper that "sucks"away the cured pattern.1 the residual liquid from the surface,leaving behind onlyA thin layer of melted wax is then spread over the surface, fillingall the voids> and cavaties left after the> removal of theresidual liq\lid. lClyer, coolingidown the waxandsl01.idif¥ingit .....Novrwe have fully solid .·.layer •. that·· is made·partlyfI"om<.

lClyer, coolingidown the waxandsl01.idif¥ingit .....Novrwe have fully solid .·.layer •. that·· is made·partlyfI"om<. plasticetnd •partly from •wax •.A cold plate is lower$donto«thesurfaceof the111The workpiece now passes under a milling disk that trims off the layer's surface down to the desired thickness, creating a flat, smooth surface ready for the next layer. A new layer of liquid photopolymer is then spread and the whole process starts once again.After the last layer is done, we have a block of wax, within which the model, or models, are embedded. microwave oven, or by using a hot air gun or even using plain warm water and the finished model is ready for use.The wax is melted away in aIf we now go back to the issue of accuracy, we can see how it is achieved in this technology. In the X and Y direction it is achieved by the full curing of the patterns, and by the highly accurate mask.

In the X and Y direction it is achieved by the full curing of the patterns, and by the highly accurate mask. In the Z direction it is achieved by the precise milling of each layer down to its desired thickness.Some other benefits that this technology offers are:Shrinkage effect has been minimized due to the full cure of every layerNo final curing in a s.pe.cialoven is neededNO··· su.pport structure is ·needed, direct.ionsthe wax supports the model in allModel$str\lct.ural strength/and stabilityare higher and •• the ·1tlodels are much less brittle. This isdu.e to/the curingpro.cessthat minimized the development of internal stresses in the structure.A.ny geometrical shape can be made without limitationHigh throughput of models within· the wax.is achieved due to the three dimensional nestingNo hazardous odorsareigenerated, for.

a very short time, and the uncured liquid is wiped off immediately.the resin stays in liquid stateThus safety is considerably higher.By altering the process slightly, wax models can be made and be used in the lost wax casting process.112•tel;1;-'-. • IitalCubital Ltd. P.O.B.330 Herzlia 46103 Israel '1"".1. 0"7?_C:':LC:?OA1C:o MASK pwrUR f} OPTICAL MASK on a FIAT GLASS PlATE e WORKPIECE CARRIAGE o HIGIl POWER IN LAMP o EXPOSURE CEIL (b RESIN APPUCATOR & AERODYNAMIC WIPER (3 MIUlNG HEAD 0) WAX APPUCATOR flO COOliNG PLATE (D AIR EXHAUSTSolider dimensions L: 4.1 ill / W: 1.7 ill /5'7" H: 1.5 ill / 4'10"Cubital America Inc. 23647 Ryan Rd. Warren, MI 48091 T~I. £212\"'IC:" "'IC:C:"'IIIIThe workstation operator creates a three dimensional composition of the models to be included in the next run and determines layer thickness. The Solider Soft ware slices the compositioninto layers and generates a precise raster image of each layer.

The Solider Soft ware slices the compositioninto layers and generates a precise raster image of each layer. Then, the following process repeats for each layer:(See highlighted components on reverse side)The raster image is sent to the mask plotter, and is converted into a high resolution optical mask on a flatThe optical mask is positioned precisely above the workpiece under the high power uv lamp, ready to "print" the new layer. • A shutter is opened for a few seconds, and the resin is cured by the light passing through the mask. The correct pattern is fully cured while the "masked" areas remain in liquid state.

• A shutter is opened for a few seconds, and the resin is cured by the light passing through the mask. The correct pattern is fully cured while the "masked" areas remain in liquid state. • The mask on the glass plate is erased and the plate is cleaned for the nextmask.The workpiece passes under an aerodynamic wiper, that collects all the uncured liquid resin, and leaves the cured areas intact.The workpiece passes under a wax ao'pncal:or melted wax is spread on top of the layer, filling all the cavities left by the liquid resin that was wiped off.A cooled plate is pressed down onto the instantly solidifying the wax, thus creating a perfect solid support.The workpiece passes under a precise .LII..aJ................ that trims the layer's surface down to the QIJ'-_'-'J.J.J.'-- ......

that trims the layer's surface down to the QIJ'-_'-'J.J.J.'-- ...... thickness and leaves a flat surface, which substrate for the next layer.The workpiece is lowered by one passes under a resin applicator thin layer of resin, ready for curing the next layer.covers it with aAfter completing the last layer, the workpiece composed of a bulk of wax within which several modelsare embedded - is removed. The wax is then either melted or rinsed and the models are ready for use.114.

oustic Analysis of Viscous Fluid Ejection Using Ultrasonic Atomizer D.A. Loney, W. Zhou, D.W. Rosen, F.L. Degertekin, A.G. Fedorov Georgia Institute of Technology 813 Ferst Drive, Atlanta, GA, 30332-0405AbstractThe acoustics of the Additive Manufacturing via Microarray Deposition (AMMD) system based on a ultrasonic atomizer is investigated for printing high viscosity fluids for 3D inkjet manufacturing applications. The ultrasonic atomizer incorporates a piezoelectric transducer, a material reservoir, and a silicon micromachined array of acoustic horn structures as ejection nozzles. When driven at the resonance frequencies of the fluid cavity, the nozzle geometry focuses the acoustic waves resulting in a locally increased pressure gradient at the nozzle apex. Previously, AMMD has demonstrated successful ejection of fluids with viscosity as high as 3000 mN-s/m2, overcoming the viscosity limitations traditionally associated with piezoelectric droplet formation.

However, the physics of ejection of such high-viscosity fluids is not well understood. This work focuses on understanding the acoustics of the AMMD system through complimentary simulations and experimental characterization. Specifically, ANSYS finite element software was used to model acoustic wave attenuation due to viscosity inside the material cavity and its implication on the pressure gradient at nozzle apex, which drives the fluid ejection. Additionally, the affect of fluid attenuation on cavity resonance modes, both the frequency and the quality factor, is characterized for fluids of a large variation range in viscosity. Finally, preliminary guidelines for improved design and efficient operation of the AMMD system are formulated based on an insight into a device’s acoustic behavior with high viscosity fluids.IntroductionInk-jet printing produces droplets from a fluid reservoir and directs the impact of those droplets onto a substrate.

Expertise developed in controlling droplet generation and impact make ink-jet processes ideal for additive manufacture (AM) applications.[1] Moreover, AM via ink-jet printing provides significant benefits over other AM techniques through scalability, low cost, and simultaneous deposition of multiple materials. Despite these benefits, AM via ink-jet printing has found limited usage due to fluid physical constraints on droplet generation, such as viscosity, surface tension, and other complex rheological behavior.[2] Of particular interest is printing of high viscosity fluids, which present a fundamental challenge of ejectability under moderate operating temperatures.Recent efforts have sought to expand the types of fluids that can be printed for AM applications beyond current viscosity limitations. Wilkes et al.

Wilkes et al. examined viscous droplet generation from a print nozzle using a capillary tube and applying a constant pressure to achieve ejection.[3] Xu and Basaran expanded this initial analysis to a sinusoidal driving force, again with capillary tube and with focus primarily on filament length as a function of viscosity.[4] In addition, slurries, powder melts, and gels have been experimentally investigated for their printability.[5-8] Despite these efforts, the choice of materials with AM continues to be limited by the ability to produce droplets either directly from the print head or from jet breakup.Figure 1: A schematic of the AMMD system, a magnified view of a single nozzle cavity, and a photograph of a device prototype.

most recently proposed a revision of the bounds upward to 4 (cid:10) (cid:4) (cid:10) 14 by defining material printability to minimize the formation of smaller, unwanted satellite droplets.[2, 11] Within systems utilizing acoustic waves for droplet generation, pressure wave attenuation due to viscous energy dissipation defines the lower boundary for printing and satellite droplet formation establishes the upper limit. As fluid attenuation is proportional to viscosity, viscosity largely determines the lower threshold on printability.

As fluid attenuation is proportional to viscosity, viscosity largely determines the lower threshold on printability. More generally, the accepted range for successful fluid printing is taken to be 1 (cid:10) (cid:4) (cid:10) 10, though these bounds are not derived based on first principles and therefore may be further adjusted with new systems and printing techniques.The Additive Manufacturing via Microarray Deposition (AMMD) system (Figure 1) consists of a piezoelectric transducer capable of generating the ultrasonic waves, a sample(1)reservoir filled with the fluid to be ejected, and an array of pyramidal nozzle micromachined in silicon acting as the acoustic wave focusing horns. When the device operates at a resonant frequency of the nozzles and fluid cavity, the acoustic horns amplify the pressure gradient locally near the nozzle orifice resulting in an efficient means for fluid ejection.

Typical device operation is in the range between 0.5 and 3.5 MHz, adjustable based on the ejector geometry, and results in either droplet-mode or jet-mode ejection based on orifice size and physical properties of the fluid.[12] Scalability is achieved by using multiple print head arrays, adjusting the number of active nozzles within each array, or by multiplexed operation of the device.[13] The AMMD system has successfully ejected fluids with viscosity as great as 3000 mN-s/m2, albeit only for short time periods at the higher range of viscosities, giving a printing indicator of approximately 0.01.[14]While the capability to eject high viscosity fluids has been demonstrated by the AMMD apparatus, the basic acoustics and fluid mechanics underlying the ejection process for such fluids has not been fully characterized. To this end, this paper focuses on acoustic behavior and examines the affect of ultrasonic pressure wave attenuation on the performance of the ultrasonic atomizer.

To this end, this paper focuses on acoustic behavior and examines the affect of ultrasonic pressure wave attenuation on the performance of the ultrasonic atomizer. This is done through a combination of numerical simulations and experiments to investigate the effect of viscous wave attenuation in the fluid on the ejection pressure gradient for different device geometries. Resulting understanding is relevant to developing guidelines for improved design and operation of the AMMD system using engineering fluids with the viscosity in the range relevant to AM applications.Model ImplementationANSYS Mechanical (ANSYS) is a commercially available, finite element software package designed to solve problems involving coupled physics.[15] ANSYS includes the capability to simulate acoustic wave propagation in a coupled fluid reservoir, nozzle plate, and piezoelectric transducer domains, allowing the full atomizer to be modeled in a single computational domain.

Pre-programmed element types provide the basis for model construction and incorporating material behavior.The piezoelectric transducer is modeled as an APC International 855 material modeled with Plane13 elements.[16] The fluid reservoir is formed by the spacer frame placed on either edge of the piezoelectric actuator and is modeled as acoustically lossless silicon having a Young’s modulus of 150 GPa, a Poisson’s ratio of 0.21, and a density of 2330 kg/m3 with element type of Plane82. The nozzle array is made of the same silicon material as the spacer and is modeled with the same element.Fluid29 is the standard fluid element for acoustic analysis within ANSYS. The mesh topology supporting Fluid29 are quadrilaterals which allow degeneration to triangles, containing two-dimensional displacement and pressure variables as degrees of freedom.

The mesh topology supporting Fluid29 are quadrilaterals which allow degeneration to triangles, containing two-dimensional displacement and pressure variables as degrees of freedom. This element includes interfacial damping at fluid structure boundaries, but assumes inviscid behavior throughout the bulk material and therefore cannot be used for viscous damping of acoustic waves.Figure 2: Nozzle geometry for the a) Fluid29 and the b) Fluid79 models. Due to the limitation on the Fluid 79 element to a rectangular shape, the nozzle slope requires division into a series of steps.Fluid79, a secondary acoustic fluid element, incorporates damping in the bulk material but presents functional limitations on implementation. Shape is constrained as rectangular with two displacement degrees of freedom. The shape constraint limits the geometry that can be meshed; any significant slope cannot be directly meshed with Fluid79.

The shape constraint limits the geometry that can be meshed; any significant slope cannot be directly meshed with Fluid79. Furthermore, any deviation from a rectangular shape rapidly introduces numerical error. Relevant material properties for this element are the bulk modulus and density, from which then ANSYS calculates the adiabatic speed of sound.Fluid79 discretizes the domain using structural mechanics formalism, as it is remapped from structural element Plane42.

The nondimensionalization can be readily obtained from common wave dispersion relations.[17] Attenuation is typically provided as a frequency (E) independent value, -B EF⁄ which grows linearly proportional to viscosity.are generally zero. ./represents aas -B C⁄ , where -BFigure 2 shows the nozzle geometry utilized with each type of fluid element. Figure 2a), meshed with Fluid29, reflects the actual geometry of the nozzle by incorporating sloped nozzle sides without approximation, as allowed by the element’s ability to conform to the slope. Figure 2b), meshed with Fluid79, utilizes a stepped approximation of the nozzle side wall geometry to allow meshing with Fluid79 elements. For full device simulations an array of nine nozzles is analyzed. It should be noted that the analysis performed is limited to 2D representation of the actual 3D shape, due to limitations imposed by the computer memory requirements for simulations.

It should be noted that the analysis performed is limited to 2D representation of the actual 3D shape, due to limitations imposed by the computer memory requirements for simulations. The pressure release boundary condition is used at the nozzle apex, and 10V AC signal is used to drive the piezoelectric transducer.To ensure the models produced similar results, the pressure distribution along the center the physically accurate (sloped), Fluid29 nozzle geometry was compared the simulation results using both Fluid29 andnozzle withtoFigure 3: Comparison of the pressure distributions along the center nozzle for the first nozzle mode of glycerol with zero attenuation and a 0.5mm spacer obtained using different ANSYS elements and domain discretization methods.Figure 4: (a) Decay in the amplitude of the pressure wave propagating along a 0.5m glycerol tube, as driven at the inlet by a harmonically-oscillating pressure of 10 Pa amplitude and having a fixed end condition.

(b) A snapshot of the waveform near the 0.1m away from the inlet of the tube. The analytical and Fluid79 simulation results essentially overlap, showing an excellent agreementFluid79 elements with the stepped nozzle for the zero attenuation case (Figure 3). A good qualitative and quantitative agreement between simulation results with the Fluid29 element for the unstepped and stepped simulation domain support the applicability of the approximate, stepped geometry with sufficiently fine steps to predict wave propagation in the nozzle horn. The comparable result between the stepped Fluid29 and Fluid79 models, particularly near the nozzle apex which is of the most interest, validates the use of the Fluid79 element. Additionally, a unit cell model for a single nozzle was developed for the Fluid79 nozzle geometry, aiming at increasing the speed of simulations.

Additionally, a unit cell model for a single nozzle was developed for the Fluid79 nozzle geometry, aiming at increasing the speed of simulations. As shown in Figure 3, the single nozzle Fluid79 simulation predicts a pressure distribution closely resembling that for the Fluid79 element simulations of the full array, thus justifying its use for computationally-demanding parametric simulations.In all simulations, convergence was assumed when further mesh refinement resulted in less than 1% change in the pressure magnitude at the central nozzle in the array. Simulations were sensitive to element size and distribution, requiring a gradual decrease in element size near the nozzle walls.

Simulations were sensitive to element size and distribution, requiring a gradual decrease in element size near the nozzle walls. Models utilizing the Fluid79 element were also intolerant of any deviation from a rectangular shape.Results and DiscussionThe capability of Fluid79 element to simulate damping behavior was first confirmed analytically by simulating a wave propagating in a one-dimensional glycerol tube, results of which are shown in Figure 4.

Relative error from analytical values was determined to be an acceptable 0.612% for the glycerol and as low as 0.053% for water under similar conditions, supporting the ability of the Fluid79 element to accurately incorporate attenuation into acoustic analysis.Figure 5: Electrical input impedance as a function of frequency for a device with 2.1 mm reservoir height loaded with glycerol and driven by a 1mm piezoelectric transducer comparing the resonant modes of the experiment and simulations.The capability of simulations using Fluid79 element to accurately predict device resonances was confirmed using the experimentally measured electrical input impedance (Agilent 8753ES Network Analyzer) of a device loaded with glycerol. The tested device had a reservoir with a height of 2.1mm and was driven by a 1 mm thick piezoelectric transducer (APC International 855). Figure 5 compares the measurements to simulation values as a function of frequency, with the nozzle resonant modes circled.

Figure 5 compares the measurements to simulation values as a function of frequency, with the nozzle resonant modes circled. Though the simulation of the nozzle array does show a slight frequency shift in the nozzle resonant frequency, most likely due to the limited number of nozzles simulated (9 as compared to 15 in the tested device), the single nozzle results closely agree with the experimentally measured resonant frequencies when normalized by the ratio of single nozzle simulation piezoelectric area to the piezoelectric area of the experimental device. The simulations also accurately capture the relative quality of different resonance modes. The quantitative differences reflect the fact that there are geometric (e.g., 2D simulations vs. 3D test device) and boundary condition differences between the model setup and a prototype package used in experiments.

3D test device) and boundary condition differences between the model setup and a prototype package used in experiments. Overall, favorable comparison between predictions and measurements provides validation of the ANSYS model developed in respect to proper coupling between the Fluid79 elements and the transducer as well as our methodology for incorporating the wave attenuation in a lossy fluid.Figures 3 and 5 demonstrate that the resonant frequencies and pressure distributions of the single nozzle simulation closely match those of the nozzle array simulation, thus supporting the use of the single nozzle model for faster, yet accurate simulations of the multi-nozzle device.

Therefore, all parametric simulations reported in the following sections have been performed using a single nozzle model of the device with the Fluid79 element.Effects of Wave Attenuation in a Lossy FluidTo assess if viscous attenuation has an effect on the fluid cavity resonance modes, shifts in nozzle resonant frequency were determined using a “test fluid” with acoustic properties ofFigure 6: Electric input impedance as a function of frequency assuming water properties but with artificially increased attenuation with a fluid reservoir and transducer thicknesses of 1mm eachFigure 7: Pressure magnitude assuming water properties with increasing damping values for a device with 2.1 mm reservoir height loaded with glycerol and driven by a 1mm piezoelectric transducerwater but with artificially increased values of attenuation for a fixed device geometry.

Attenuation was varied from - EF⁄ (cid:5) 30 H 10I>J per meter, corresponding to water, to - EF⁄ (cid:5) 30000 H 10I>Jper meter, an order of magnitude greater than glycerol. As Figure 6 demonstrates, the resonant frequencies of the device appear to be independent of the magnitude of the attenuation parameter.Figure 7 shows the standing pressure wave distribution along the centerline of the fluid reservoir and the nozzle cavity, as a function of the attenuation coefficient. The results indicate that attenuation parameter smaller than - EF⁄ (cid:5) 30000 H 10I>J per meter causes only small pressure amplitude reductions due to the short path that a wave travels within the device fluidFigure 8: Pressure gradient at the nozzle apex as a function of resonant frequency, assuming water properties and increasing values of attenuation. cavity.

cavity. For higher values of -B EF⁄ , pressure amplitude reduction becomes significant but focusing behavior of the acoustic horn formed by the nozzle structure remains present with an amplified pressure gradient at the nozzle apex near fluid ejection point.The driving mechanism for fluid ejection is the pressure gradient generated at the nozzle apex, shown in Figure 8 as a function of the attenuation parameter. Although there is a significant difference in the pressure gradient among different resonant modes of the device, for low values of -B EF⁄ , the pressure gradient for each resonant mode changes little, owing to a short path for wave propagation within the cavity. As the attenuation parameter increases, the pressure gradient becomes reduced for all modes, except the first, fundamental cavity mode (~0.7 MHz) and the mode (~2.1 MHz) closest to the piezoelectric transducer’s resonance (~2.2 MHz).

It is interesting to note that modes whose frequency is lower than the transducer resonance generally yield a greater pressure gradient than modes whose frequency is greater than the transducer resonance.Effects of Device GeometryFigures 9 and 10 display the effect of changing the fluid reservoir height on the pressure gradient at the nozzle apex for glycerol and water, respectively. As expected, the magnitude of the pressure gradient is much greater across all the modes for a less lossy fluid (water) as compared to more viscous glycerol. Glycerol also shows a distinct trend of increased pressure gradients for resonant modes occurring near the piezoelectric transducer resonance.

Glycerol also shows a distinct trend of increased pressure gradients for resonant modes occurring near the piezoelectric transducer resonance. In contrast, water displays a more uniform pressure gradient across all resonant modes, with an elevated pressure gradient for certain modes with no particular dependence on the location with respect toFigure 9: Pressure gradient at the nozzle apex as a function of resonant frequency for various fluid reservoir heights, assuming glycerol properties (K LM⁄ (cid:5) NOOO H POIPQ per meter) and a 1mm thick transducerFigure 10: Pressure gradient at the nozzle apex as a function of resonant frequency for various fluid reservoir heights, assuming water properties (K LM⁄ (cid:5) MQ H POIPQ per meter) and a 1mm thick transducerFigure 11: Pressure gradient at the nozzle apex as a function of resonant frequency for various transducer thicknesses, assuming glycerol properties and a reservoir height of 1mmthe transducer resonance.

The magnitude of the pressure gradient for both fluids is largely independent of the fluid reservoir height and can be targeted by selecting an ejection mode at a fixed fluid height. Moreover, the dependence of the glycerol pressure gradient on the transducer resonant frequency implies one may vary both ejection mode and the transducer height/material to obtain an optimal pressure gradient for ejection. However, the lack of such dependence in water indicates that device dimension changes may have limited effect for low viscosity fluids where the pressure wave attenuation due to attenuation is less between resonant modes.Figure 10 further examines the dependence of the pressure gradient at the nozzle apex on the transducer resonant frequency. The transducer resonant frequency is adjusted by increasingthe thickness of the material, from 0.5 mm to 2 mm, lowering its resonant frequency.

The transducer resonant frequency is adjusted by increasingthe thickness of the material, from 0.5 mm to 2 mm, lowering its resonant frequency. In general, as the transducer resonant frequency shifts, the fluid cavity modes nearest the transducer frequency always yield the largest pressure gradients. However, those cavity modes near the transducer frequency do not attain equal magnitude for all transducer frequencies. For example, the third cavity mode at ~2 MHz frequency for 1 mm thick fluid reservoir driven by a 1 mm piezoelectric transducer yields a pressure gradient almost twice as large as all other cases. This leads to the conclusion than an optimal piezoelectric transducer resonance frequency exists for a given fluid and fluid reservoir height such as to produce a maximum pressure gradient at the ejection point.

Performing such an optimization would be critical for printing high viscosity fluids where maximizing the pressure gradient is of the paramount importance to achieve ejectability.ConclusionDroplet generation from high viscosity (low printing indicator) fluids continues to be a challenge for AM applications. The AMMD system based on an ultrasonic atomizer has shown the ability to print fluids with significantly greater viscosity that traditional ink-jet based methods, but optimization of device design for stable, sustained operation requires an in-depth understanding of physical fields underlying fluid ejection. This paper focuses on the development of simulation tools and the prediction of an acoustic field in the ultrasonic atomizer with a particular emphasis placed on accounting for the pressure wave attenuation by highly viscous fluids. The standard ANSYS Fluid79 element has been shown to be capable of incorporating attenuation into an acoustic analysis.

The standard ANSYS Fluid79 element has been shown to be capable of incorporating attenuation into an acoustic analysis. Through the incorporation of attenuation, it has been shown that increased attenuation does not affect the AMMD nozzle resonant frequency, given that other fluid properties are held constant. Furthermore, for varying fluid cavity heights of water and glycerol – two model fluids having properties representative of many materials of interest to AM applications – the pressure gradient has been shown to be largely independent of fluid attenuations values for a fixed ejector geometry.

Finally, it has been shown that by adjusting the fluid reservoir height and the piezoelectric transducer resonant frequency, one can optimize the apparatus to yield a maximum in the pressure gradient, facilitating the printing of traditionally unprintable materials.AcknowledgementsWe would like to express our appreciation to the National Science Foundation for award DMI-0900322 which supports our work. Any opinions, findings, and conclusions or recommendations expressed in this publication are those of the authors and do not necessarily reflect the views of the National Science Foundation. We would also like to express our gratitude to Dr. J. Mark Meacham for his assistance with simulations and experiments discussed in this paper.Bibliography 1.Calvert, P., Inkjet printing for materials and devices. Chemistry of Materials, 2001. 13(10): p. 3299-3305. Jang, D., D. Kim, and J. Moon, Influence of Fluid Physical Properties on Ink-Jet Printability. Langmuir, 2009. 25(5): p.

Unstable geometrical characteristics of single tracks and high roughness of the powder layer can cause porosity in 3 dimensional printed parts. In this study a non-destructive online monitoring technique, using acoustic emission was utilized to determine lack of fusion and balling effect of single tracks. This phenomenon was simulated by using an increased powder layer thickness. Short Time Fourier Transform was used as a tool for analysis of the acoustic behaviour of the system and it was compared with the acoustic emission (AE) recorded during processing of single tracks.2087IntroductionLaser based powder bed fusion (LPBF) is an additive manufacturing (AM) process in which thermal energy from a laser beam selectively fuses regions of a powder bed [1].

The benefits of LPBF are: high degree of design optimization derived from shape complexity; possibility to produce complex lattice and combined structures; relatively high resolution, limited by laser beam size and powder-based properties; re-usability of powder material; etc.

But due to the track-by-track, layer-by-layer nature, LPBF has drawbacks such as: bad surface finishing, size and design limitations - especially for overhanging parts; high thermal gradients which cause residual stress that can result in the formation of internal cracks and deformations during processing; non-optimal input process parameters such as scanning and building strategies or inhomogeneity in delivered powder layers cause porosity in LPBF parts; defects cause a substantial deterioration in the performance properties in LPBF parts and limit wide application of this Additive manufacturing (AM) technology in the industry [2].In-process defect sensing and control is one of the main steps for ensuring repeatability and consistency of LPBF manufacturing. Key process parameters in LPBF are: laser and scanning parameters; powder material properties; powder bed properties and recoating parameters and build environment.

Key process parameters in LPBF are: laser and scanning parameters; powder material properties; powder bed properties and recoating parameters and build environment. Process signatures emanating from the melt pools emission/absorption; radiation;reflected/scattered light, etc.-these phenomena are the basis to control stability and repeatability of the LPBF process [3]. Visual, ultrasonic, Eddy current, radiographic, magnetic methods, liquid penetrant test, shearography, acoustic and thermography – are all NDT technique that can be used for quality control in AM [4], [5]. Thermography and acoustic emission testing are applied as inspection methods for in-situ monitoring of AM processes, but the spatial and temporal resolution of in-process monitoring with feedback depends on specific scanning parameters. LPBF is a very fast process, so on-line NDT and feedback control is extremely difficult to attain.

LPBF is a very fast process, so on-line NDT and feedback control is extremely difficult to attain. For spatial capability, minimum AM defect size also has to be investigated in order to determine if the part is fit for service [6-7]. Available commercial LPBF monitoring systems mainly use: photodiode melt pool monitoring; CMOS camera for powder bed imaging system; as well as others methods that include the use of IR and UV photo sensors and pyrometers [3].in LPBFare: molten/solidifiedpool;plasmaAn in-process non-contact method could be AE - AE is seen to have high signal to noise ratio and fast response. Non-contact methods can be used for quality control of the ever growing laser technology implemented in industrial applications [8]. Palanco, et al. (2003)[9] studied microphone signals to monitor laser ablation and acoustic emission of plasmas . It was shown that spectral analysis of acoustic waves is a reliable technique for the diagnostics of laser plasma phenomena.

It was shown that spectral analysis of acoustic waves is a reliable technique for the diagnostics of laser plasma phenomena. Mao et al. [8] 1993 showed that the acoustic spectrum of conduction welding is different from that of keyhole welding and that there exists strong correlation between the AE energy signal and laser power, welding speed and focusing distance (laser spot size).For laser sintering, surface modifications and welding, various phenomena such as: melting and solidification; vaporization; interaction of materials with protective atmosphere; shrinkage and solid-state phase transformations; plastic deformations and cracking - these2088are all related to acoustic phenomena. In general, qualitative types of AE signals are burst (discrete) and continuous. Delamination, crack initiation under deformation or corrosion can be qualified as energy “burst” that is emitted.

In general, qualitative types of AE signals are burst (discrete) and continuous. Delamination, crack initiation under deformation or corrosion can be qualified as energy “burst” that is emitted. Acoustic signals from diffusive phase transformations or coalescence of micro-cracks that can be classified as continuous signals [7], [10], [11]. AE in materials that undergo deformations and fractures depends on physical properties of material as well as environmental factors [12], [13]. AE testing is exceptionally sensitive when compared to other NDT techniques and can recognize crack growth of the order of 25 μm [7].Leaks and frictions i.e. interactions of media in relative motion, chemical reactions and changes of size of magnetic domains also generate acoustic waves and create other class of AE signals that can be studied for quality control of tooling and manufacturing. Charde et al.

Charde et al. [14] (2016) interpreted the weld formation using AE for the carbon steels and stainless steels welds in servo-based resistance spot welding. A typical acoustic behaviour of servo- based system was found and then AE amplitude of frequencies and corresponding phase shifts for “ideal” welding condition was received at expulsion welding conditions were analysed. It was found that different AE patterns were used to distinguish the behaviour of ideal and faulty welding processes that were qualified by macrographs. Similar results were received for laser welding of Al and Polyamide sheets by Schiry et al. [15] (2016). Rough joint was a result of excessive energy input. Gas bubbles provoked creation of the gas channels registered by hits-time AE monitoring during the process. Higher hits in the area of the gas channels correlated with the appearance of the holes in welding.

Higher hits in the area of the gas channels correlated with the appearance of the holes in welding. With decreased energy input, weak joint with low pulling strength was created and AE signals were irregular and wide scattering was found. Also it was shown that the material as well as distance of the sensor to the weld spot has a strong influence to the signal.A non-contact acoustic inspection method for LPBF was proposed by Redding et al.[16] (2017) - it is opposed to sensors being attached underneath the build substrate. Signal profiles for free-defect parts were compared with manufactured, qualified parts. Sudden deviations in the amplitude of acoustic signals indicated a fault in the process. A similar approach was used by Gaja and Liou [17] (2017) for defect classification of AM by direct energy deposition process - amplitude, ringdown count, duration, rise time, i.e. time dependent AE event features were collected and used for clustering analysis.

time dependent AE event features were collected and used for clustering analysis. Cracks and pores and their AE signal characteristics were successfully distinguished and verified by post-test optical microscopy. Fisher et al. [18] (2016) showed that a notable difference could be obtained from the acoustic signatures of varying laser powers during LPBF. It was found that there was a clear shift and missing peaks in spectral analysis between the two different laser powers.AE has great potential for monitoring machine events during AM processing [4]. Combination of AE with machine learning for in-situ and real-time quality monitoring of the AM process was reported by Wasmer et al.[19] (2017) and Shevchik et al. [20] (2018). The sensor was mounted directly inside the process chamber and it was used to detect AE for in- situ quality monitoring of LPBF process. Signals were analysed by spectral and conventional convolutional neural networks.

Signals were analysed by spectral and conventional convolutional neural networks. It was found that a correlation between AE signals and2089porosity of the samples can be made, however it was also shown that AE from the AM process was quite weak compared to a strong noisy background.A method of analysis for AE could be a Short Time Fourier Transform (STFT) - STFT is used to obtain time-frequency analysis of data signals. STFT calculates the energy distribution in the joint time-frequency domain. STFT is accomplished by calculating the Fast Fourier Transform (FFT) of the signal in a sliding window, this window is moved across the signal to give a representation of the frequency content at that specific portion in time [21].The goal of the study is to determine if AE can be used to detect lack of fusion and balling phenomena.

Scanning on an increased layer thickness was compared to scanning on optimal powder layer thickness in order to establish the relationship between process parameters of laser scanning of maraging steel powder and AE signatures from LPBF by EOSINT M280 system (EOS, GmbH).Materials and methodsAs stated the purpose of this paper to identify a non-destructive acoustic emission technique to determine lack of fusion and balling-effect of single tracks. Firstly, balling phenomenon was simulated by increasing the powder layer thickness to 300 µm. A microphone was placed inside the chamber to optimally sample data. The next step was to put the acquired data set through a predetermined high pass filer to allow for analysis of the acoustic behaviour of the system. Comparison of AE was obtained utilizing STFT as a tool.

Comparison of AE was obtained utilizing STFT as a tool. Finally, conclusions are drawn on indicators and frequency identifiers of defects from the results of the STFT.The experiment was done using Maraging steel with chemical composition: Fe-18Ni- Co-5Mo-1Ti being Ni 17.6%, Co 8.88%, Mo 4.85, Ti 1.06%. Samples were produced on the substrate with similar chemical composition. The building chamber was filled with nitrogen atmosphere. Three single tracks, 200 mm in length were scanned at a laser power of 305 W with a scanning speed of 1.1 m/s with an EOSINT M280 system. AE was measured using an ICP microphone having an optimal frequency range of 3.75-20 000 Hz (±2dB). The microphone was placed inside the building chamber (Figure 1).2090Figure 1: Schematic of the experimentThe data was acquired at a sampling frequency of 102.5 kHz.

The microphone was placed inside the building chamber (Figure 1).2090Figure 1: Schematic of the experimentThe data was acquired at a sampling frequency of 102.5 kHz. It was seen experimentally that to remove the effect of ambient operating noise that does not pertain to the actual laser scanning a high pass filter could be used. The optimal low cut off frequency was seen to be 1500 Hz. This was done by sampling a section of the recording before the tracks were scanned, applying an STFT and analysing the results.After filtering the data signal consisted of the laser and possible AE defects. Gaincontrol and a Short time Fourier transform (STFT) were applied to obtain the time domain as a tool to analyze and evaluate the relevant defect detections indicators.Results and discussionAs stated, analysis, data gathering and transform processing with STFT was accomplished by track morphology and acoustic emission.

Thus the result of the STFT evaluations and analysis will be used to determine markers for porosity forming phenomena, in other words fusion defects.The result of the first step of simulated fusion and balling can be seen in the Figure 2 below, which shows the top view of single tracks at optimal and non-optimal scanning parameters.A5mmmBAB10mmFigure 2: Maraging steel tracks on the substrate at 305 W laser power and scanning speed of 1.1m/s at 50 µm powder layer thickness (Optimal; A) and 300 µm (non-optimal; B)2091The top view of the optimal process parameters show that the track is uniform and continuous, while the non-optimal shows that the track is clearly irregular and almost no track formation is present. The top two images in Figure 2 show control tracks at higher magnification. Note the three control scan tracks and the three defect scan tracks.

The top two images in Figure 2 show control tracks at higher magnification. Note the three control scan tracks and the three defect scan tracks. Using STFT the three control tracks will be analyzed against the three defect scan tracks.Acoustic emissionThe figure below shows the results of the next step which is the high pass filter which is needed prior to analysis of recordings to remove unwanted environmental noises and improving the relevant signal to noise ratio.The entire signal of the process from start to finish is shown in Figure 3.

It can be seen that the EOSINT M280 has various processes that emits sound, but no clear scanning is shown until the filtering and processing is applied.Figure 3: Graph of entire process over time optimal (left) and non-optimal(right)The three tracks can be clearly distinguished using the high pass filter as shown inFigure 4, the time of scanning correlates to that of scanning a 200mm track at 1.1 m/s.Figure 4: Signal of three tracks at optimal (left) and non-optimal (right) before(top) and after (bottom) applying signal filtering2092The next step is the application of the STFT to analyze and evaluate the scan samples. Note in figure 5 that when using a Short Time Fourier Trasnform (STFT) one can clearly distinguish when scanning starts and when each track was being scanned. Figure 5 below shows a spectrogram of the three tracks in both optimal and non-optiaml cases.

Figure 5 below shows a spectrogram of the three tracks in both optimal and non-optiaml cases. The spectogram is a three dimensional graph with colour indicating amplitude with blue as low values and red as high values. The y-axis is the corresponding frequencies and the x-axis the time interval of the sampled data. Each time interval on the x-axis will show the Fourier transform in a window of 102 samples and show a colour indicating the magnitude of that specific frequency in the y-axis.Figure 5: Spectrograms of three scans showing a decrease in intensity in non-optimal processing, optimal (left) and non-optimal (right)In figure 5 the three different tracks can be clearly seen in the time domain - again these scanning times correlate to that of scanning a 200mm track at 1.1m/s.

The frequency content of the scans is spread over the entire range of frequencies, with high frequencies being more prominent in the non-optimal scan.The next step in the analysis is to add all the values in the individual frequencies across the x-axis together. This should give an indication as to which frequencies has the most energy or significance, with the low values of the between scans not influencing the result. This is done by subtracting this result of the control (normal layer) with the result of the thick layer (that consist of lack of fusion and balling effects). Thereafter results should give an indication on possible identifiers for defects. The results are seen below in Figure 6.2093Figure 6: Energy content of optimal and non-optimal scanning over frequency range (left). The energy difference between optimal and non-optimal scanning (right)In Figure 6, the y-axis indicates the magnitude and the x-axis indicates the different frequencies.

The energy difference between optimal and non-optimal scanning (right)In Figure 6, the y-axis indicates the magnitude and the x-axis indicates the different frequencies. The left graph shows the magnitude of the two different signals, white indicating the optimal tracks and red indicating the non-optimal tracks. The right graph shows the difference in the magnitude of the optimal and non-optimal signals. It can be seen that the results of the spectrogram in Figure 5 is more prominently shown in Figure 6. If there was no difference in the optimal vs non-optimal then the result in Figure 6 (right) would have been zero. This frequency graph shows large variances. The right side of Figure 6 shows that the large indicators are present in high frequency range.

This frequency graph shows large variances. The right side of Figure 6 shows that the large indicators are present in high frequency range. The differences can be used as an indicator for online detection of porosity forming phenomena during fusion or lack of fusion in metal laser powder bed fusion.ConclusionThe purpose of this paper was to identify an online monitoring AE technique to determine lack of fusion of single tracks. Much research could still be done, for instance: identifying the minimum layer thickness for error detection using AE; ideal AE pre-filtering and the effect of different materials. The significance of this paper is indicators for online detection of porosity forming phenomena such as lack of fusion during metal LPBF. This study shows that STFT could be a valuable tool in the analysis of online detection of porosity forming phenomena during metal LPBF. An automated online process can be developed using these defect indicators.

An automated online process can be developed using these defect indicators. To conclude these nondestructive AE identifiers could be used to validate metal LPBF parts.AcknowledgmentThis work is based on the research supported by the South African Research Chairs Initiative of the Department of Science and Technology and National Research Foundation of South Africa (Grant №97994) and the Collaborative Program in Additive Manufacturing (Contract №CSIR-NLC-CPAM-15-MOA-CUT-01). The authors are grateful to IMCE for carrying out the impulse excitation measurements.2094References1.2.3.4.5.6.7.ISO 17296:2015. Additive manufacturing –General principles – Part 2: Overview of process categories and feedstock. ISO 2015, Geneva. Quinlan, H. E., Hasan, T., Jaddou, J. and Hart, A. J. 2017. Industrial and consumer uses of additive manufacturing: A discussion of capabilities, trajectories, and challenges. Journal of industrial ecology, 21: S15-S20. Spears, T.G. and Gold, S.A.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperACTIVE - Z PRINTING: A NEW APPROACH TO INCREASING3D PRINTED PART STRENGTHJivtesh B. Khurana*, Shantanab Dinda†, and Timothy W. Simpson*† *Department of Mechanical and Nuclear Engineering †Department of Industrial and Manufacturing Engineering The Pennsylvania State University, University Park, PA, 16802AbstractResearch suggests that topology and build parameters in Fused Filament Fabrication (FFF) play a vital role in determining mechanical properties of parts produced by this technique. In particular, the use of 2D layers printed parallel to the build surface produces high anisotropy in parts making them the weakest when loaded perpendicular to the layer interfaces.

In particular, the use of 2D layers printed parallel to the build surface produces high anisotropy in parts making them the weakest when loaded perpendicular to the layer interfaces. We investigate a novel approach that uses non-planar 3D layer shapes - Active Z printing, to improve mechanical strength through alignment of localized stress tensors parallel to the deposition paths. Sinusoidal layer shapes are used with varying amplitude, frequency, and orientation. Design of experiments is performed to correlate effect of varying shape and orientation of sinusoidal layer shapes on flexural strength of parts. Based on this, the results are used to decide parameters to be studied further and characterize their effect on the strength of parts.1. IntroductionAdditive Manufacturing (AM) is the process of creating end use parts by adding material in a layer by layer manner.

IntroductionAdditive Manufacturing (AM) is the process of creating end use parts by adding material in a layer by layer manner. Fused Filament Fabrication (FFF) is a process, where a thermoplastic is melted into liquid state in an extruder head and subsequently deposited through a nozzle. This nozzle traces the part’s cross-sectional geometry to produce 3D parts directly from CAD data in a layer by layer manner1. The layers used in this process are conventionally two dimensional, and printed parallel to the build surface. As a consequence of this layer by layer approach, polymer parts produced by FFF are anisotropic, making them especially weak at the interface between adjacent layers and deposition toolpaths1,2,3.Fused Filament Fabrication (FFF) using 2D layersFigure 1: Fused Filament Fabrication Processhave been extensively investigated to improve the mechanical properties through optimization of process parameters1.

In the past, researchers have explored the idea of non-planar deposition paths through the study of Curved Layer FFF (CLFFF). Active-Z includes the use of dynamic tool paths that contain Z movements within individual layers, as compared to conventional FFF toolpath that have constant Z values in intra-layer extrusions.1627Active – Z printing uses distinct layer shapes oriented relative to a part in order to produce improvement in flexural properties of parts produced using AM. An open source slicer, Bread, was used which accepts part geometry and layer shape as an STL file and produced Gcode that can be interpreted by a 3D printing machine. The Gcode produced is further post processed in order to remove unwanted travel moves, nozzle to part collisions and optimize travel moves to reduce print time. We used a Rostock Max V2 Delta 3D printer due to its capability to produce smooth and rapid Z movements.

We used a Rostock Max V2 Delta 3D printer due to its capability to produce smooth and rapid Z movements. The machine interface was performed using Printrun Pronterface software which is also open-source. All flexural specimens were produced in the XYZ orientation with infill toolpath orientations alternating 0o/90o from the Y -axis of the machine. Flexural testing was performed following the procedure outlined in ASTM D790 and data was collected for the experimental design.2. Literature Review2.1 Modeling and effect of Curved Layer Fused Filament FabricationCurved Layer Fused Filament Fabrication (CLFFF) was introduced by Chakraborty et al. to improve mechanical properties, and reduce stair stepping and print time. Their approach uses offsetting of part surfaces using Bezier surfaces to create 3D layers7, however, no physical tests were conducted. Singamneni et al.

Their approach uses offsetting of part surfaces using Bezier surfaces to create 3D layers7, however, no physical tests were conducted. Singamneni et al. proposed another technique that used generation of surface point data from STL files and then offsetting points in the Z direction, followed by taking a cross product of 2 vectors at every surface point4. Parts produced by this technique were subjected to flexural loading, resulting in a 40% increase in load before fracture when compared with conventional parts showing potential for strength improvement using CFFF. Huang et al. studied the effect of varying raster angle on the flexural properties and concluded that strength decreases with increasing raster angle, however, this relation is non-linear5. They also developed an adaptive slicing method to vary layer height for CLFFF, tests showed improved resolution of fine and sharply varying surface profiles as well as direct relation between flexural strength and layer thickness8.

Furthermore, Yerazunis et al. used hemispherical-shaped specimens to test CLFFF parts in shear and bending, and concluded that parts produced using CLFFF are 3.5 to 5 times the strength of conventionally printed parts6.Another major application of CLFFF has been towards recreating skins and surfaces of parts with small curvature. Conventional Flat Layer Fused Filament Fabrication (FLFFF) produces stair-stepping and inferior mechanical properties, this can be easily rectified using CLFFF. Llewellyn-Jones et al. conducted case studies using shell like geometries of shoe in-sole, dished sandwich panel and a car body panel9. Using these case studies, they showed that it was possible to combine curved layer and conventional statics Z layers to produce superior surface finish and minimize risk of part failure from layer delamination. Similar studies were conducted by Allen et al. on using curved toolpath for skin of parts.

Similar studies were conducted by Allen et al. on using curved toolpath for skin of parts. Innovative slicing and printing methods were implemented and an improvement in supported surface quality was observed. They suggested that these techniques provide improvements in mechanical properties, however, physical studies have not been conducted10. An analogous process was implemented for concrete panels by Lim1628et al. They improve surface finish by minimizing staircase effect, reduced printing time by minimizing layer count, and increased filament bonding by minimizing voids and maximizing the contact area between layers11.2.2 AM Process - Property – Structure RelationshipsIdentification of Process-Property-Structure are imperative for every manufacturing process; especially in AM, where anisotropic properties can emerge due to factors such as material deposition pattern, build orientation and material interface.

It is known that processing parameters affect fiber to fiber bond strength and mesostructure of parts. Furthermore, uneven thermal cycles due to layered deposition leads to accumulation of stresses, causing distortion which is primarily responsible for poor interlayer bonding, subsequently reducing strength (Section 2.1).Existing research has shown that processing parameters play an important role in providing optimum conditions for strength development in FFF processed parts. These are raster angle, raster width, layer thickness, air gap, part orientation, contour width, and extrusion & envelope temperatures. Es-Said et al. showed that molecules align themselves in the direction of raster deposition causing decrease in resistance to flexural load, hence, decreasing flexural strength with increasing raster angle13. Similar experiments for raster angle conducted by Constance et al. found the same relation, however, failure surface showed shear failure at large raster angles.

Similar experiments for raster angle conducted by Constance et al. found the same relation, however, failure surface showed shear failure at large raster angles. These effects contribute to increased resistance to load due to long raster lengths associated with small raster angles14. Ahn et al. found that raster angle and air gap significantly contributed to tensile and flexural strength as compared to raster width, model temperature, and filament color using a fractional factorial experimental design1. Moreover, Sood et al. varied layer thickness, build orientation, raster angle, raster width, and air gap using a central composite experiment design and analyzed their relationships to tensile, compressive, and flexural strength using response surface methodology. They found that reduction in distortion and increase in inter-extrusion diffusion caused increase in strength.

They found that reduction in distortion and increase in inter-extrusion diffusion caused increase in strength. These effects were primarily seen with large raster widths, large number of layers produced as a result of combination of layer heights and part orientation15, 16. Further, small values of air gap and raster angle favored high strength similar to the results by Ahn et al. and Es-Said et al.; Wang et al. optimized flexural modulus for complaint bow structure using a L9 orthogonal taguchi array. It was found that air gap influenced the modulus the most followed by raster angle, raster width, and slice height17.Apart from working parameters, machine parameters such as extrusion temperature, envelope temperature, nozzle diameter, and feed-rate also play a vital role in determining strength of parts.

As the extruded plastic cools, internal stresses are developed due to uneven deposition speeds resulting in layer deformation in the form of warping, cracking, and layer de-lamination. These phenomena greatly affect part strength and dimensional accuracy18. Moreover, it was determined that these deformations increase with increasing model height. However, these deformities can be eradicated by increasing chamber temperature to the glass transition temperature of the deposition material. Bellehumeur et al. used the polymer sintering model to evaluate the effects of extrusion and envelope temperatures, and dimensions of the extruded1629filaments. Extrusion temperature plays a vital role in neck growth of the bonding zone as temperature cannot be maintained for long durations for bonding between filaments to complete. As a result, mechanical properties of the bonding zone are inferior to virgin ABS material19.Bellehumeur et al.

As a result, mechanical properties of the bonding zone are inferior to virgin ABS material19.Bellehumeur et al. further found that temperature at bottom layers rises above the glass transition temperature and rapidly decreases in the direction of movement of extrusion head. This temperature increases with the number of layers; hence bottom layer extrusions experience higher diffusion as compared to the rest of the part20. Joseph et al. used polymer rheology data and polymer weld theory to predict strength of parts subject to uniaxial tension. It was found that inter- extrusion bond was fully healed. Based on these results, the strength was predicted to be slightly over the 95% confidence level of experimental strength21.2.3 Research ScopeAlthough, CLFFF has been modeled and demonstrated, the applications of non-planar sinusoidal layers are yet to be modeled and studied.

Active-Z printing provides the ability to define sinusoidal layer geometry in order to precisely vary material properties inside a part. Given that mechanical properties of parts are closely linked to layer shape and orientation, achieving an apt understanding of how part properties relate to process parameters is paramount to ensure its relevance and reliability as an AM process.In this paper, the authors seek to probe the process-property-structure relationships of Active-Z printing through a design of experiments. A fractional factorial design is used to analyze the effect of five parameters on the flexural strength and modulus of specimens, (i) Amplitude/Pitch ratio (layer shape), (ii) Z-Rotational angle, (iii) X-Rotational Angle, (iv) layer height, (v) and air gap.

The research question is “Do non-planar sinusoidal layer shapes cause an increase in the flexural strength and modulus of parts produced using Active-Z printing? If so which factors significantly affect them?” Active-Z Printing is explained in detail in section 3. The experimental design is outlined in section 4 followed by results and discussions in section 5 & 6 respectively.3. Active-Z PrintingActive-Z printing is a branch of FFF technique which involves the use of motion in X, Y, Z axis simultaneously to produce shapes with 3D cross-sectional area. This is contrary to conventional FFF which only uses planar layers. Active-Z printing can be divided into two design tracks, first, the design of desired part and second, design of layers. The design of layers allows designers to effectively control the material behavior of the part in addition to the geometry. A process workflow for using Active-Z is shown in figure 2.

The design of layers allows designers to effectively control the material behavior of the part in addition to the geometry. A process workflow for using Active-Z is shown in figure 2. In this study, a sinusoidal herringbone structure is used as a layer shapes. These herringbone structures were inspired by bio-structures investigated by Yaraghi et al. who found sinusoidal herringbone structures in dactyl club of a mantis shrimp25. These geometries were found to have increased young’s modulus and fracture strains as compared flat helicoidal structures.1630Figure 2 : Design workflow for Active-Z PrintingIn order to slice layers in such a shape an open source slicer, Bread23, is used which accepts (i) the part to be sliced as an STL, (ii) the shape of the layer as an STL, (iii) and a configuration file highlighting the process parameters used to manufacture the part. The slicer exports toolpath as G-code which can be readily accepted by AM machines.

The slicer exports toolpath as G-code which can be readily accepted by AM machines. However, the slicing software is under development and is prone to producing a range of unwanted travel moves which lengthen print time as well as cause collisions between the nozzle and the specimen to be printed. A MATLAB script was used to remove unwanted travel moves, check for nozzle-part collision and create fast and optimum travel path.4. Experimental Design and Methods4.1 Parameter SelectionMechanical properties of FFF parts are closely linked to processing parameters like orientation, extrusion width, raster angle, layer height, and air gap. Further, strength development between individual roads is highly dependent on the extrusion temperature and scan speed (Section 2.2). With the induction of Active-Z printing the authors feel that the part’s mechanical properties might be affected by layer’s shape and orientation.

With the induction of Active-Z printing the authors feel that the part’s mechanical properties might be affected by layer’s shape and orientation. While raster angle has been shown to affect CLFFF parts, the layer shape and orientation has not yet been studied.1631A fractional factorial test has been used by to characterize the effect of layer shape and orientation in the Active Z process. A cause and affect diagram (aka Ishikawa diagram) was created to group factors that affect strength of 3D printed parts and is shown in figure 3.

A cause and affect diagram (aka Ishikawa diagram) was created to group factors that affect strength of 3D printed parts and is shown in figure 3. The purpose of cause and affect diagrams is to systematically group causes of specific event into families leading to identification of the family containing the root cause.the authorsFigure 3: Cause and effect diagram from FFFAlthough all branches of the Ishikawa Diagram show the factors that could affect the material properties of the specimen, the “Layer Properties” and “Process Parameter” branch was further analyzed due to its uniqueness to Active-Z. Figure 4 depicts the 5 parameters chosen for analysis.

Figure 4 depicts the 5 parameters chosen for analysis. The description and hypothesized effect for each parameter is shown in Table 1.Table 1: List of parameters used for design of experimentsParameterAir gapAmplitude/pitchZ- Rotation angleX- Rotation angleDescription: In-plane distance between two adjacent extrusions in a cross section of a layer Hypothesis: Varying air gap directly effects part density and thus strength decreases with increasing air gap. Description: A measure of the shape of sinusoid, a large ratio signifies a thin and stretched out shape while a small ratio signifies fat and bulged shape. Hypothesis: The sinusoids act like springs; the ratio would be directly proportional to stiffness of part. Further, load applied is resolved into a normal and shear load due to varying shape of part. Description: It is the angle between the x-axis of layer and the x-axis of the part measured counter clockwise from the part’s x-axis.

Description: It is the angle between the x-axis of layer and the x-axis of the part measured counter clockwise from the part’s x-axis. Hypothesis: Small angles would result a larger road length increasing resistance to load. Description: It is the angle between the normal vector of layer shape and z- axis of the part measured clockwise from the part’s z-axis. Hypothesis: Increase in angle would result in a further distribution of load along layer normal and layer parallel vectors, thus increasing strength.Layer heightDescription: The thickness of each layer used to build the part, it is often an integral multiple of the size of part. Hypothesis: It is know that flexural strength increases with increasing layer height15, interaction with layer shape would cause increased strength.1632cabdeFigure 4: List of parameters tested.

(a) Z Rotation Angle (b) Layer Height (c) Air gap (d) X Rotation Angle (e) Amplitude/pitch4.2 Design of ExperimentsThe goal of the experiment is to identify parameters that affect the flexural strength and modulus of parts created using Active-Z Printing. Theories introduced in Design and Analysis of Experiments24 are consulted to produce a 5 parameter 2 level fraction factorial design of experiments (DOE). A linear relationship between the response and the five parameters was desired, hence, only 2 levels one set high (+1), one set low (-1) was chosen. Appropriate level of each parameter was decided based on nozzle shape, machine limitations and literature review.A half factorial experimental design was devised for the study. A defining relation of E = ABCD was chosen and higher order effects were confounded with main and second level interaction effects. This allows us to run 16 distinct tests with each test replicated thrice to quantify standard error within each experiment.

This allows us to run 16 distinct tests with each test replicated thrice to quantify standard error within each experiment. As a result, a total of 48 test specimen were used, making the DOE relatively easy to regulate and execute. A list of coded parameters in the factorial designed experiment is shown in table 2.Table 2: Process parameters tested and their respective testing levelsSymbol A B C D EParameter Air Gap Amplitude/Pitch Ratio X-Rotational angle Z-Rotational angle Layer HeightUnit mm N/A Degree Degree MmLevel -1 1.25 0.09858 30 60 0.3Level 1 3 0.14997 60 75 0.416334.3 Experimental ProcedureASTM D790-15e2 was followed to measure specimens’ flexural strength and modulus. This standard has been used in engineering design and allows comparison of material properties across manufacturing platforms, including existing AM processes. Figure 5 summarizes dimensions for the selected flexural test specimen.

Figure 5 summarizes dimensions for the selected flexural test specimen. A length of 96mm was used with a thickness of 6mm and a width of 12mm. This allows for an 80mm span with 8mm overhang and a span to depth ratio of 13.33:1. Specimens’ lengths and widths were measured to monitor any variability in size. Due to the experimental nature of the slicing algorithm and the nozzle geometry, it was found that the specimen’s dimensions were 12.96 ± 0.504 mm wide and 6.03 ± 0.145 mm thick. Similarly, the specimens’ masses were measured on a digital scale to monitor to variation in mass, found to be ± 1g.Each experiment (table 3) was printed in the run order using a SeeMeCNC Rostock Max V2. Parts were printed at a uniform speed of 30mm/s with an extrusion temperature of 210 ± 2º C and a bed temperature of 105 ± 1 º C, as recommended by the filament manufacturer. All temperature measurements were taken using a 100K EPCOS thermistors rated for temperature measurement up to 300º C.

All temperature measurements were taken using a 100K EPCOS thermistors rated for temperature measurement up to 300º C. Each specimen was printed individually, a total of 48 build trays were printed in this study with each specimen in the XYZ orientation. Following the completed print job, the build plate was allowed to cool down from 105 ± 2º C to 65 ± 2º C before the specimens were removed to minimize warping. Upon removal, the specimens were stored in an airtight transparent bag in a metal cabinet located away from potential sources of moisture.Tensile tests were conducted using an MTS 800 material testing machine in accordance to ASTM D790 standard. Using TestWorks for data acquisition with the machine, tests were conducted at a stable environment of 25 º C, 25 - 50% humidity and tested at a crosshead federate of 1.78 mm/min until specimens ruptured on the tensile load side.

Specimen deflection was measured using crosshead position in accordance to the D790 Standard testing procedure type I.Figure 5: Size of the test specimen used for flexural testingFigure 6: Printed specimen showing sinusoidal layer shape16345. Results and Statistical Analysis5.1 Experimental DataThe results were obtained from flexural testing of all 16 specimens with 3 replicates each. The mean and standard deviation of each experiment is summarized in table 3. To perform statistical analysis on the data obtained, the assumptions made for the analysis of the factorial design are first verified. A normal probability of the runs was calculated for Flexural Strength and modulus. This follows a linear trend, indicating that the data is an acceptable normal probability distribution with a confidence of 96.4% and 97.1% for flexural strength and modulus respectively.

This follows a linear trend, indicating that the data is an acceptable normal probability distribution with a confidence of 96.4% and 97.1% for flexural strength and modulus respectively. Further, an outlier analysis was performed on the dataset to find datasets that lies an abnormally large distance away from other random samples in a population. For the flexural strength, no outliers exist, however, the flexural modulus does show an outlier (Experiment Id 3) but it can be ignored given there were replicates for each specimen. This outlier would add to the pure error of the model.Finally, a residual analysis was conducted for both flexural strength and modulus. Residuals for each portray a normal distribution, versus fits and versus order plots show a balanced response and the run order does not show any time sequence.

(MPa) 0.40 0.71 6.72 0.03 2.42 0.66 0.69 1.45 0.85 0.37 0.41 1.35 0.26 0.35 0.24 0.53Flexural Modulus (MPa) 1117 1012 1939 1482 1110 1425 1348 1330 1593 1154 1434 1514 1482 666 1534 1103Standard Dev. (MPa) 105.5 56.4 135.8 38.9 118.6 135.3 47.1 16.3 84.1 43.2 55.1 77.9 40.5 51.0 6.2 45.85.2 Statistical AnalysisUsing the experimental data and neglecting aforementioned parameters, a factorial analysis was performed again to determine the effects of a parameter and its interactions on flexural strength and modulus. The resolution V factorial design used will be able to estimate all main effects and all two factor interactions. Thus, the model would produce 16 terms, the intercept term, the 5 main effects, and the 10 two-factor interactions.The factorial model fitted shows a R2 and R2 (adj) value of 95.74% and 93.75% respectively for flexural strength, and 95.25 and 93.03% respectively for flexural modulus, implying a good fit.

Normal plots are used to graphically convey which main and higher order effects are statistically significant.Normal Plot of the Standardized Effects (response is FLM, Alpha = 0.05)Normal Plot of the Standardized Effects (response is Stress, Alpha = 0.05)999995DEffect Ty pe Not Significant Significant95Dt n e c r e P90807060 504030ADBEDEE ABBDF actor N ame A B C D EA B C D Et n e c r e P908070 60 50 4030ADDECEBEBCAB20CCE20ACBE10AC10C5A5A1a115105Standardized Effect0515100 Standardized Effect55Figure 9: Normal probability plot for (a) flexural modulus and (b) flexural strength1636Effect Ty pe Not Significant SignificantF actor N ame A B C D EA B C D EbEffects from a consistent process would form a normal distribution, thence statistically significant effects would act as outliers for the dataset.

In the plots below statistically significant effects are shown in red boxes whereas insignificant effects are portrayed as black dots.For flexural strength, the largest absolute effects were air gap and X rotation angle. For second level interactions, the largest absolute effects were air gap*X rotation angle and amplitude/pitch*layer height. Further, all main parameters, i.e., Air Gap, Amplitude/Pitch ratio, Z Rotation angle, X Rotation angle, and Layer Height were found to be statistically significant. Flexural strength increases with increasing Z angle whereas inverse is true for all other parameters. Air gap*Layer Height (A*E), Amplitude/Pitch*Z Rotation angle (B*D), and X Rotation angle* Z Rotation angle (C*D) interactions were statistically insignificant, hence, were removed from further analysis. Similar to flexural strength, for flexural modulus, the largest absolute effects were Air Gap (A) and X rotation angle (C).

Similar to flexural strength, for flexural modulus, the largest absolute effects were Air Gap (A) and X rotation angle (C). While, Air gap*X Rotation angle (A*C) had the largest effect for a second level interaction. All main effects except Amplitude/Pitch were statistically significant. Air gap*Layer height (A*E), Amplitude/Pitch *X Rotation angle (B*C), and X Rotation angle* Z Rotation angle (C*D) were found to be statistically insignificant and were neglected from further analysis. However, given Amplitude/Pitch*Layer Thickness (BE) and amplitude/pitch*Z rotation angle are statistically significant for flexural modulus, Amplitude/Pitch cannot be neglect as a parameter.

ANOVA is a statistical technique where the sum of squares and error terms are compared against an F-ratio statistic along with p-values to determine the statistical significance of the parameters. The DOE was modeled using least squared sums where the P-value was calculated for each parameter and compared against critical P ratio of 0.05 (95% confidence window) calculated.

Interaction effect plots were created to validate interactions and if they showed significant effects towards response in the model, figure 10 & 11. Interaction plot show how the relationship between one categorical factor and a continuous response depends on the value of the second categorical factor. Parallel lines indicate no interactions while non-parallel lines show interactions.1638Interactions can be divided into ordinal (non-intersection) and disordinal (crossover) interactions. Ordinal interactions have a difference in the mean of each factor for the interaction indicating a main effect of contributing factors as well as an interaction between them. On the other hand, disordinal interactions signify that the mean effect of factor must be qualified in terms of the impact of the second variable and mean effect cannot be interpreted independently.

Amplitude/Pitch* Z angle (BD), Amplitude/Pitch*Layer thickness (BE) on the other hand cross over and are classified as disordinal interactions. Based on these interactions we can conclude that for flexural modulus, main effects for amplitude/pitch and X rotation angle and layer height cannot be interpreted individually. The final effects individually significant for flexural strength and flexural modulus are summarized in table 8.Table 8: Summary of significant effects for flexural stress and flexural modulusMain Effects2nd order interactionsEffectsAir gap (A)Amplitude Pitch Ratio (B)X Angle (C)Z Angle (D)Layer height (E)AB AC AD BC BD BE CE DEFlexural StrengthYESNOYESYESNOYES YES YES YES NO YES YES YESFlexural ModulusYESNOYESNONOYES YES YES NO YES YES YES YES16396. Discussions6.1 Analysis of ResultsFrom table 5 and figure 10 it can be noted that flexural strength increases for decreasing air gap (A).

Discussions6.1 Analysis of ResultsFrom table 5 and figure 10 it can be noted that flexural strength increases for decreasing air gap (A). Air gap is responsible for the decrease in part density due to increased spacing between adjacent roads. Less material is present to bear load, hence, strength decreases. Z rotational angle (D) defines the orientation of high and low sections of a sine wave in the part. This orientation of high and low section of the sine wave have an effect similar to that of raster angle. As raster angle increases the raster length and resistance to bending load decreases however distortion due to thermal gradients also decrease18. In case of small rater lengths, the reduction of raster distortion dominates causing an increase in strength15. X rotation angle (C) plays an important role in determining the magnitude of shear load acting in the transverse direction from the x-axis of the layer plane.

X rotation angle (C) plays an important role in determining the magnitude of shear load acting in the transverse direction from the x-axis of the layer plane. As X rotation angle increases the shear load decreases, load normal to layer increases and strength decreases. This theory is further backed up by the fact fracture surfaces show tensile failure for large X rotation angles.The amplitude/pitch ratio (B) and layer thickness (E) demonstrate a significant interaction. At higher values of layer thickness, as amplitude/pitch ratio increases the sine function becomes thin and elongated therefore, the number of fully formed sine wavelets in specimen increases resulting in an increase in strength. However, at lower values of layer height the distortion effect coupled with reduced number of fully formed sine waves associated with large amplitude/pitch ratio results in a decrease in strength.

However, at lower values of layer height the distortion effect coupled with reduced number of fully formed sine waves associated with large amplitude/pitch ratio results in a decrease in strength. But, the increase in diffusion15 at small layer thickness dominates the interaction between Z rotation angle (D) and layer thickness (E) increasing strength for large values of Z rotation angle and decreasing layer height. It is to be noted that for interaction between air gap (A) and X rotation angle (C), the effect is a combination of increase in density related with small rater angles and large shear forces associated with large X angles resulting in increased strength.From interaction plots shown in Figure 11, it can be observed that the effect for air gap (A), X rotation angle (C) and Z rotation angle (D) on flexural modulus is similar to what is observed in the case of flexural strength.

However, the interaction effects of air gap*amplitude/pitch ratio (AB), air gap*X rotation angle (AC), amplitude/pitch ratio*layer thickness (B*E) and amplitude/pitch ratio*Z rotation angle (BD) show a different behavior. It is known that modulus of stress increases with increasing amplitude/pitch ratio25, this effect coupled with increased contact area between adjacent rows and layers and full sine wavelets at lower layer heights result in increasing stiffness. Further, increased amplitude/pitch (B) coupled with smaller loads normal to the layer for large Z rotation angle (D) leads to increased flexural modulus. In the case of air gap*amplitude/pitch ratio (A*B) interaction, the effect of amplitude/pitch assisted with increased part density due to small air gap lead to an increase in stiffness of part. Finally, for interaction of air gap*X rotation angle (A*C), increased raster length dominates distortion effects to produce increasing stiffness for small air gap.

Finally, for interaction of air gap*X rotation angle (A*C), increased raster length dominates distortion effects to produce increasing stiffness for small air gap. Based on the analysis presented maximum flexural strength is expected to be found at A-1B-1C-1D1E1 and max flexural modulus at A-16401B1C1D1E-1. The experimental run for maximum observed flexural modulus was found to coincide with the predicted max flexural strength.6.3 Sources of Experimental ErrorThe design of experiment used assumes that the response is purely linear and accounts for variation from the linear trend, however, apart from the given choice and magnitude of parameters used, the response is not likely to be linear. Even though the model used has a good statistical fit (p-value above 90%), replicating experiments with more than 2 levels would be beneficial for an adept understanding of the Active-Z process.

Even though the model used has a good statistical fit (p-value above 90%), replicating experiments with more than 2 levels would be beneficial for an adept understanding of the Active-Z process. Apart from the statistical analysis, the MTS 81- material testing machine used a load cell of 500kN which produced large amounts of noise for small loads used in the testing. A low pass filter was used to minimize noise; however, a more adequate load cell would be better suited. During the test flexural test, crosshead displacement was used to measure maximum central displacement in the specimen which is prone to sag in mechanical linkages. Further, changes in temperature and relative humidity between experiments and testing were not explicitly controlled, affecting the hygroscopic nature of ABS polymers used.A large variation of size was observed in the as manufactured specimens, in particular for parts with small layer thickness, small Z rotation angle and large X rotation angles.

These variation in width can be attributed due to the nozzle geometry rubbing against the part deviating it from the programmed path.Regression analysis of the main effects produced an R2 value of 59%. This indicates that either the data is not as linear as expected, or there are parameters that have not been considered. The former is being currently verified with a 3-level design, but requires a significant number of experimental parts and tests. In the case of the latter, additional parameters seen in figure 3 are considered responsible. Though not included for this study, a similar analysis of such additional parameters will be performed in the future. Therefore, this experimental analysis is considered to be part of a larger study to verify factors affecting performance in Active-Z printing and quantify these effects.7.

Therefore, this experimental analysis is considered to be part of a larger study to verify factors affecting performance in Active-Z printing and quantify these effects.7. ClosureThe purpose of the study was to investigate the effect of sinusoidal layers on the mechanical properties of parts printed with Active-Z printing and isolate process parameters to study further. The authors employed design of experiments to identify critical factors affecting flexural strength and modulus. The air gap, layer shape, and orientation about the part and layer thickness were varied and statistical analysis found that air gap, X rotation angle and amplitude/pitch ratio were the most significant effect on the tested material properties.

Small air gap, small X rotation angles constituted towards stronger parts, while specimens with large amplitude/pitch ratio were found to be stiffer.1641This paper provides the framework for future work towards understand effect of repeating sinusoidal structures and their orientation on strength of parts. The authors will study significant parameters to further characterize the shape of a response, and develop a model assist application of this technique towards optimization of mechanical parts. Apart from process characterization, there is a need for a robust slicing algorithm that can effectively calculate layer to part intersection, and produce toolpath that takes strength development in material welds into consideration. An AM process which can vary mechanical properties locally inside a part would add to a designer’s ability to develop stronger optimized parts for end use.8. AcknowledgementsThe authors acknowledge Dr.

AcknowledgementsThe authors acknowledge Dr. Charles Bakis, Professor of Engineering Sciences and Mechanics at Penn State, for providing access to the MTS 800 for material testing. The authors also acknowledge Joseph Bartolai, PH.D. candidate at Penn State’s Center of Innovative Materials Process using Direct Digital Deposition (CIMP-3D), for his assistance with flexural testing. Finally, this project was funded in part by the Pennsylvania State University’s College of Engineering Research Initiative (CERI) and Erikson Summer Discovery Grant.16429. References1. Ahn, Sung‐Hoon, M ontero, M ., Odell, D., Roundy, S., & W right, P. K “Anisotropic Material Properties of Fused Deposition Modeling ABS.” Rapid Prototyping Journal, vol. 8, no. 4, Oct. 2002, pp. 248–257.2. Gibson, I., Rosen, D. W. & Stucker, B. Additive Manufacturing Technologies: Rapid Prototyping to Direct Digital Manufacturing, Springer, New York, 2010.3. ISO/ASTM.

TIVE DEVICE FABRICATION USING FIBER ENCAPSULATIONADDITIVE MANUFACTURINGM. Saari, M. Galla, B. Cox, E. Richer, P. Krueger, and A. Cohen*Laboratory for Additive Manufacturing, Robotics, and Automation (LAMRA), Department of Mechanical Engineering, Bobby B. Lyle School of Engineering, Southern Methodist University, Dallas, TX 75275Corresponding authorAbstractFiber Encapsulation Additive Manufacturing (FEAM) is a novel solid freeform fabrication process in which a fiber and a matrix are co-deposited simultaneously within a single printer along straight and curved 2-D and 3-D paths. Using a FEAM approach in which the fiber is a metal wire and the matrix is a thermoplastic polymer, simple electromechanical devices such as voice coils, inductive sensors, and membrane switches have been successfully produced.

This paper will present an overview of the FEAM process, describe several fabricated devices, and discuss recent developments in controllably stopping and starting the wire, and in creating electrical junctions between individual wires, which together enable much more complex devices to be made.Introduction and MotivationIn recent years a number of researchers have investigated the additive manufacturing (AM) of multi-material, multi-functional structures and active devices. These efforts, which attempt to go well beyond the purely structural and aesthetic dimensions of single-material AM, seek to achieve a higher level of functionality and performance, or significant reductions in size, weight, and cost, for a variety of electromechanical and electronic systems.

Such benefits, combined with AM’s intrinsic ability to rapidly produce complex, customized geometries, promise a future in which previously-impossible designs of great sophistication can readily be produced, a future in which AM is truly enabling. An area of interest for the present work, and for which multi-material AM holds great promise, is soft robotics.Soft robotics has emerged as a research topic of considerable interest [Majidi 2013, Luand Kim, 2013], in part driven by the desire to create “co-robots” that work safely in close proximity to humans. Moreover, robots with soft bodies and limbs might grasp and manipulate delicate and irregular objects—such as picking fruit off of a tree without damage. They might also move in new and complex ways, and transform their shapes: capabilities that open novel application areas, such as search and rescue operations.

They might also move in new and complex ways, and transform their shapes: capabilities that open novel application areas, such as search and rescue operations. The tentacle of an octopus, if it could be emulated artificially, is an example of a device useful in a soft robot that would enable complexmotions and shapes. Yet in contemplating how to implement such a device, one realizes there are a very large number of degrees of freedom making it difficult to implement by traditional manufacturing techniques. In some robot designs, one might also wish to provide redundancy of actuators so as to increase reliability and robustness. In both of these cases, there is a fundamental need to provide multiple, distributed actuators, which are not normally found in conventional, rigid robots. Distributed actuators are likely to need closed-loop control, which implies a distributed array of sensors, e.g., one per actuator.

Distributed actuators are likely to need closed-loop control, which implies a distributed array of sensors, e.g., one per actuator. Moreover, it can be beneficial to sense the outside world on the entire robot “skin” with reasonable spatial resolution (i.e., an array of tactile sensors); this also requires a distributed approach. Thus, soft robots can achieve their full potential only if their bodies and limbs can be equipped with distributed network of actuators and sensors, and interconnects that allow power and signals to be routed between them.Requirements and ApproachThe problem then becomes how to practically manufacture such a system. Assembling robots with tens to hundreds of interconnected actuators and sensors from discrete components would clearly be costly and labor-intensive, and would likely result in a device of considerable volume and weight (e.g., due to the redundant packages of individual components).

This leads to the question: might it be possible to directly, monolithically manufacture functional robot components, and therefore bypass these issues? More specifically, might it be feasible to use a multi-material AM process to literally print a robot limb in which all the required actuators, sensors, and wiring are built-in, with no assembly required, such that the limb is essentially ready for use? Indeed, if this could be achieved, the potential applications would extend considerably beyond use in soft robotics.Monolithic fabrication of macroscale devices through additive manufacturing clearlyrequires working with both dielectrics having good structural properties and with good electrical conductors. Specifically, electromechanical devices, as well as electronic circuits, depend on conductors for their operation, such as connecting them to signals or power sources, or generating magnetic fields (e.g., in a solenoid actuator or motor).

A number of activities in printed electronics—both 2-D and 3-D—have been reported. However, printing good conductors in combination with dielectrics is very challenging. We define a “good conductor”—one suitable for real-world applications with reasonably high currents—as meeting the following requirements:Low electrical resistivity and high maximum current density.  Flexibility, strength, and ductility, ideally similar to a soft metal, allowing the conductor to be strained or absorb impact energy while in use without fracture.  Thermal stability, so as to remain functional at elevated temperatures.Ability to be easily and reliably interfaced to other materials and components.  Low total cost, so that the conductors do not make the printed device unaffordable for all but a few applications.Printing of the conductor along with dielectric material introduces additional requirements.

At a minimum, the process should be compatible with the processing of the dielectric material, produce a unified structure when completed (so that the two components are well-integrated), offer large flexibility in the device design, and operate at reasonable manufacturing speeds.Though excellent progress has been made, the most common approach used in printedelectronics, namely printable inks, still faces significant challenges. Rather than striving to develop yet another conductive material that might meet the requirements, we have taken a different approach. We acknowledge that the best “good conductor” is—and is likely to remain for the foreseeable future—solid metal wire. Wire has very low resistivity (e.g., 1.7 µohm-cm for copper), allows high current density (e.g., 20 A/mm2 for chassis wiring, which is important for electromagnetic actuators), is well-characterized electrically, strong, and depending on composition, can be flexible and ductile (e.g., copper).

It is thermally stable and is easily interfaced through well-established methods such as soldering, welding, and thermosonic bonding. Moreover, conductive wire is inexpensive, especially compared with silver-based inks, and available in a variety of forms useful for building sensors and actuators. Overall, the favorable characteristics of conductive wires far outpace printable conductive materials compatible with thermoplastics.Other Research using WireSome previous work has considered incorporating solid conductive wires into additivemanufacturing systems. A project known as “SpoolHead” aimed at integrating solid conductors directly into printed parts was undertaken several years ago [Bayless et al., 2010]. A system was developed to place wire on top of a printed layer of thermoplastic material, bond it using a heated tip, and cut it. Some simple demonstrations were achieved, but the system had difficulty in bonding the wire securely to the polymer surface and was slow.

Some simple demonstrations were achieved, but the system had difficulty in bonding the wire securely to the polymer surface and was slow. Additionally, no attempt was made to join wires to one another. More recently, researchers have leveraged earlier work on ultrasonic wire embedding for the production of smart card antennas [Taban, 2003], and applied it to embedding wire into 3-D printed thermoplastic objects. Wire is embedded below the surface by ultrasonic vibration or heating which causes the thermoplastic to reflow [Aguilera et al., 2013]. A workcell has been constructed comprising a six-axis robot that transfers partially-build parts from station to station in a temperature-controlled enclosure [Abriz et al., 2014]. The workcell includes two FDM systems, and apparatus for wire embedding, laser welding (to createjunctions), and a CNC mill.

The workcell includes two FDM systems, and apparatus for wire embedding, laser welding (to createjunctions), and a CNC mill. This system has been shown to work, but involves costly capital equipment, has much slower processing than typical AM systems (e.g., due to the time required to move parts between stations), and requires that the part be made from a thermoplastic material, so it can reflow.Fiber Encapsulation Additive ManufacturingHaving selected metal wire as the conductor in our strategy to monolithically print macroscale electromechanical devices, the goal is then to develop an AM process that can print both wire and dielectric material, and do so using a system that is reasonably fast, easily automated, integrated (e.g., a single machine), and affordable. The process we are developing, Fiber Encapsulation Additive Manufacturing (FEAM), has the potential to meet this goal.

The process we are developing, Fiber Encapsulation Additive Manufacturing (FEAM), has the potential to meet this goal. FEAM has similarities to material extrusion AM processes such as Fused Deposition Modeling (FDM), but provides for the simultaneous, co-deposition and encapsulation of a fiber (metal wire, carbon fiber, Kevlar®, optical fiber, etc.) along with a solidifiable, fluid matrix (dielectric polymer such as molten thermoplastic or thermoset resin, ceramic, etc.) [Saari, Cox, Richer et al., 2015]. As shown in Fig. 1(left), both the fiber and the matrix (e.g., copper wire and thermoplastic resin) form a “coaxial composite” in which the fiber is fully encapsulated within the solidified extrudate of matrix material.The key challenges in making FEAM a robust process having the versatility to produce awide range of electromechanical and electronic devices are:1. Keeping the wire substantially centered within the extrudate, even when followingtoolpaths with small radii.2.

Keeping the wire substantially centered within the extrudate, even when followingtoolpaths with small radii.2. Controllably starting and stopping the wire, so that wire can be placed where required.3. Creating reliable junctions between wires with reasonably low resistance: both inter-layer (between layers) and intra-layer (within a single layer).Fig. 1: (left) A coaxial composite; (right) schematic of the FEAM printhead (right).Progress on addressing the first of these challenges has been discussed [Saari, Cox, Richer et al., 2015]. In this paper, we report on initial results addressing the latter two challenges, and describe several early proof-of-concept demonstrations of FEAM’s capability and potential.The FEAM SystemIn FEAM, both the fiber (e.g., copper wire) and the matrix (e.g., a thermoplastic polymer such as ABS) are deposited using a single printhead.

The printhead comprises an extruder with a nozzle similar to those used in FDM, and a guide positioned to the side of the nozzle orifice, through which the wire is delivered (Fig. 1, right). The axis of the guide is approximately horizontal, allowing the wire to be introduced into the polymer as it is extruded from the orifice while the printhead moves. The printhead must follow a toolpath that is generally curved according to the contours of the layer. Because it is asymmetric, (and the wire must stay within the extrudate and be reasonably well centered), the guide must rotate around the nozzle orifice, remaining approximately tangent to the nozzle velocity. Alternatively, the guide may remain fixed relative to the FEAM system, and the part may rotate around the nozzle. Since the latter approach provides more flexibility and facilitates material handling, it was selected for the FEAM testbed shown in Fig. 2 (left).

Since the latter approach provides more flexibility and facilitates material handling, it was selected for the FEAM testbed shown in Fig. 2 (left). The testbed incorporates a rotational, or “theta” stage, onto which X and Y stages, which translate the part, are mounted. With the orifice aligned to the theta rotational axis, rotation of the theta stage results in the guide rotating around the nozzle with respect to the part. Fig. 2 (left) also shows the Z stage of the machine and a filament extruder with a nozzle. Not shown in the left-hand figure, but shown in the right-hand closeup, is the wire feeder/cutter mechanism which is mounted to the Z stage, adjacent to the nozzle.Fig. 2: (left) FEAM testbed; (right) detail of wire feeder/cutter.Wire may be fed through the guide passively (i.e., once anchored in the solidifiedextrudate, it can be pulled through the guide) or actively. Active feeding, however, is required for starting the wire after it has been stopped (i.e., cut).

Active feeding, however, is required for starting the wire after it has been stopped (i.e., cut). The feeder/cutter (Fig. 2, right) includes a stepper motor to feed the wire through the guide, and a solenoid-operated blade to cut the wire. Cutting is performed “upstream” of the guide and nozzle, with each wire segment cut and then pushed through the guide by the advancing wire behind it.Addressing the ChallengesWire encapsulation and centering. Initial FEAM experiments focused on producing straight extrudates with nickel and copper wire encapsulated in BendLay Tough [Amazon, 2015], a transparent ABS-like material which allowed the wire position within the extrudate to be more easily evaluated. Results of this were promising, with the wire remaining reasonably centered, so curved toolpaths were evaluated next.

Results of this were promising, with the wire remaining reasonably centered, so curved toolpaths were evaluated next. By programming the testbed controller to simultaneously move the X, Y, and Z stages to follow a helical toolpath while rotating the theta stage to maintain tangency, “thin” vertical coils were printed from continuous nickel or copper wire (76-127 µm diameter) insulated with BendLay Tough. Fig. 3(left) depicts a fabricated cylindrical coil having an outside diameter (O.D.) of 25 mm and a helix pitch of 250 µm. Several smaller coils are described in an earlier publication [Saari, Cox, Richer et al., 2015].It was found that when printing smaller diameter coils, the wire would not typically remain centered horizontally and in some cases exited the extrudate entirely.

This was addressed in part by reducing the printing speed, suggesting that the molten extrudate required some time to solidify in order to “lock” the wire in place before radial forces on the wire could displace it significantly. In addition, when cooling air was directed by a small nozzle onto the molten extrudate the solidification was observed to accelerate dramatically. No apparent reduction in mutual adhesion of extruded polymer was observed when cooling air was used, and even tall coils (e.g., 150 mm) could be built without any problems. Initial, non-optimized printing speeds that were tested with cooling air have been described earlier 1 [Saari, Cox, Richer et al., 2015], and ranged from 3.2-25.4 mm/s. Very small turning radii could be achieved. Square coils (Fig. 3 (right)) were printed with air cooling and the corner radius of curvature of the wire was measured to be approximately 600 µm [Saari, Cox, Richer et al., 2015].

Square coils (Fig. 3 (right)) were printed with air cooling and the corner radius of curvature of the wire was measured to be approximately 600 µm [Saari, Cox, Richer et al., 2015]. Testing was later done to determine whether the largest coil could be printed at a higher speed. Indeed, stable printing at a speed of nearly 50 mm/s—comparable to standard, single-material FDM coil—was achieved, and cooling for this diameter coil was not required.One of the fabricated coils was metallographically mounted, sectioned, and polished, asdescribed earlier [Saari, Cox, Richer et al., 2015]. Horizontal (i.e., X/Y) centering of the wire1 Table 1 in that publication has an error in the smallest O.D. shown; the correct value is 2.8 mmwas good, though there were small variations in both extrudate and wire position. Vertical centering was acceptable, though the wire was slightly lower than desired. In more recent tests we have shown that this can be improved.Fig. 3.

Vertical centering was acceptable, though the wire was slightly lower than desired. In more recent tests we have shown that this can be improved.Fig. 3. (Left) cylindrical copper coil approximately 25 mm in outside diameter; (right) square copper coilapproximately 25 mm wide.Stopping and starting the wire. Stopping and starting the flow of the liquid matrix (e.g., extruded thermoplastic polymer) is straightforward, and is common practice in FDM. Doing the same with wire is far more difficult. In the current approach, the wire is cut with a solenoid- activated blade. If the blade is sharp, it cuts with minimal burr, and can be automatically re-fed by upstream drive rollers without jamming. Using the feeder cutter of Fig. 2 (right), test parts such as the “dashed line” part shown in the backlit views of Fig. 5 were produced with 100% feeder/cutter reliability over the duration of the test.

2 (right), test parts such as the “dashed line” part shown in the backlit views of Fig. 5 were produced with 100% feeder/cutter reliability over the duration of the test. In this part, segments of 127-µm copper wire with a programmed length of 3, 13, and 25-mm were fed, cut, and encapsulated—with space between the segments—within BendLay Tough extrudates forming a portion of a slab printed in the same material. In the top frame of the figure, the majority of the slab can be seen, while the lower left view is a closeup showing magnified wire segments. Finally, the lower right view shows the highly-magnified end of a segment. Unfortunately, due to refraction of light through the polymer extrudates, it is difficult to obtain a clearer view of the wire segments.Printing junctions.

Unfortunately, due to refraction of light through the polymer extrudates, it is difficult to obtain a clearer view of the wire segments.Printing junctions. Coils produced as described above, in which the wire is continuous and the coil is a 3-D shape may useful as discrete devices, but in the context of 3-D printing a more complex system comprising, for example, a group of electromagnetic actuators, sensors, and interconnects, a standard layer-by-layer approach is required. In other words, coils should preferably be made from discrete curved segments of wire which are cut on each layer—freeing up the printhead to deposit material elsewhere—and then joined electrically at a junction, as inFig. 6 (left), where the ends of the wire (here shown with a square cross-section) terminate in inter-layer junctions that allow the electrical pathway to be continued between layers.Fig. 5: Fed and cut segments of copper wire encapsulated within a clear polymer part.

5: Fed and cut segments of copper wire encapsulated within a clear polymer part. (top) overview;(bottom) closeup views.To produce junctions between wires, a number of methods may be used, including laserwelding, soldering, and thermosonic bonding. We have chosen to focus on yet another approach, which does not depend on close contact or precise alignment between wires, and has the benefits of speed and relative simplicity. While silver-based inks and related materials have a number of liabilities such as high resistivity, high cost, and questionable mechanical properties, these issues are greatly mitigated if the use of these materials is limited to junctions between wires, not the wires themselves. Indeed, a resistivity as high as 10,000 µohm-cm may be acceptable in such an application, since the distances between wires are short (e.g., 150 µm).

Indeed, a resistivity as high as 10,000 µohm-cm may be acceptable in such an application, since the distances between wires are short (e.g., 150 µm). Moreover, materials used in junctions are consumed in small quantities, minimizing cost, and because they are short mechanically, they are less fragile. Given this, we are developing a silver particle-based electrically conductive polymer composite (ECPC) intended for junctions. Such a material canbe used to encapsulate one or more neighboring wires in an intra-layer junction, shorting them together as in Fig. 6 (right, top). Or, regions of ECPC surrounding wires on adjacent layers will be in electrical contact, shorting the wires together to form an inter-layer junction (Fig. 6 (right, bottom)).Fig. 6: (left) Model of layered coil with junctions; (right); intra-layer and inter-layer junctions.ECPCs can be formulated with different polymers, including both thermoplastics andthermosets.

6: (left) Model of layered coil with junctions; (right); intra-layer and inter-layer junctions.ECPCs can be formulated with different polymers, including both thermoplastics andthermosets. An ECPC based on a thermoplastic polymer, however, can be printed using a suitable heated extruder, much like pure thermoplastics are printed in FDM. The key requirements are that the ECPC components (particulate and binder material) do not segregate in the extruder and that the material does not clog a small nozzle and remains sufficiently conductive as-deposited.Fig. 7. (left): Intra-layer junction test part; (right) closeup of ECPC junction.The current ECPC formulation is based on silver-coated nickel powder and athermoplastic elastomer. It has been measured to have a typical static resistivity of approximately7,000 µohm-cm, though values as low as 3,000 µohm -cm have been observed.

It has been measured to have a typical static resistivity of approximately7,000 µohm-cm, though values as low as 3,000 µohm -cm have been observed. It can also be extruded through the 0.7 mm nozzle of a mini screw-based extruder [Saari, Cox, Galla et al., 2015], though we believe some segregation of elastomer and particles may be occurring, and good control of flow initiation and termination requires further development. Fig. 7 shows an overview (left) and a closeup (right) of an intra-layer junction test part in which two wires having a programmed horizontal separation of 0.38 mm (center-to-center) were encapsulated in BendLay Tough except in a central cavity open to the top surface. ECPC with a resistivity of 10,000 µohm-cm was deposited into the cavity by the extruder, where it rapidly solidified. The resistance between the wires was then measured using a model RM3544-01 resistance meter (Hioki, Ueda, Japan). The best measured resistance of several samples was 250 mohm.Fig.

The resistance between the wires was then measured using a model RM3544-01 resistance meter (Hioki, Ueda, Japan). The best measured resistance of several samples was 250 mohm.Fig. 8 shows a closeup (left) of an inter-layer junction test part in which two wires separated vertically by 0.25 mm (center-to-center) were encapsulated in BendLay Tough except in a central cavity on each of the two layers. The same ECPC was deposited into each cavity, such that the ECPC deposits overlapped. The best measured resistance of this junction was 110 mohm. A COMSOL simulation of the junction geometry was performed (Fig. 8 (right)) using the nominal wire geometry and the measured bulk resistivity of the ECPC. This yielded a predicted junction resistance of 40 mohm. We believe the higher measured value may be due to voids in the ECPC surrounding the wires, however, this has not been verified.Fig.

This yielded a predicted junction resistance of 40 mohm. We believe the higher measured value may be due to voids in the ECPC surrounding the wires, however, this has not been verified.Fig. 8 (left): Closeup of inter-layer junction; (right) COMSOL simulation results.Demonstration DevicesA number of thin cylindrical coil inductors such as that shown in Fig. 3 were fabricated and characterized, and several devices based on thin coils were demonstrated and described, as were membrane switches using two crossed wires [Saari, Cox, Richer et al., 2015]. Coil-based devices included an inductive eddy current sensor, a linear variable differential transformer(LVDT), a rheostat, and a loudspeaker.

Coil-based devices included an inductive eddy current sensor, a linear variable differential transformer(LVDT), a rheostat, and a loudspeaker. A capacitive force sensor built monolithically from thermoplastic elastomer, BendLay Tough, and copper was also constructed and characterized [Saari, Cox, Galla et al., 2015].More recently, FEAM has been used to print “thick” coils, in which there are multiple spiral turns of wire on each layer. Thick coils were made with continuous 127 µm copper wire printed in a spiral pattern with 380 µm pitch, without need for junctions. Printing was alternated layer-by-layer between outside-in spirals or inside-out spirals. Fig. 9 (left) shows such a coil, measuring 25 mm O.D. and 20 mm tall. The coil has 800 turns of wire and an approximate resistance and inductance of 54 ohms and 4300 µH at 10 KHz, respectively.

Fig. 9 (left) shows such a coil, measuring 25 mm O.D. and 20 mm tall. The coil has 800 turns of wire and an approximate resistance and inductance of 54 ohms and 4300 µH at 10 KHz, respectively. The coil was assembled along with a plunger support having ABS flexures, and a frame and plunger printed from magnetic iron PLA (ProtoPlant, Vancouver, WA), to form a solenoid. An iron frame was also made by machining, and a steel plunger was taken from a commercial solenoid.To test the pull force, solenoids with different frames and plungers were inverted so theplungers rested on a laboratory balance and extended approximately 1 mm from their fully- retracted positions. The weight loss produced by briefly energizing the solenoid with a nominal current of 0.5 A at a voltage of 31V was used as a measure of solenoid pull force. Results are shown in Table 1 for various frame and plunger materials. Clearly, iron-PLA did not perform well, especially as a plunger.

Results are shown in Table 1 for various frame and plunger materials. Clearly, iron-PLA did not perform well, especially as a plunger. This is not surprising since, based on density measurements, this material has an iron content of approximately 10%, making it only weakly magnetic. Also, it can be seen that while use of iron or steel for the frame instead of iron-PLA improves the pull force significantly, the improvement is far less significant than substituting steel for iron-PLA in the plunger.Fig. 9. (left) Multi-turn “thick” coil inside a cutaway iron-PLA frame; (right) electric bell incorporatingcoil, complete frame, ball-tipped moving plunger, and metal gong.The solenoid was also used to create an electric bell as shown in Fig. 9 (right) by placing the flexure-supported plunger beneath a steel gong, and driving it using a square wave oscillatorat 30V peak-peak at frequencies of 15-20 Hz.

9 (right) by placing the flexure-supported plunger beneath a steel gong, and driving it using a square wave oscillatorat 30V peak-peak at frequencies of 15-20 Hz. Two configurations were tested: an iron-PLA frame and plunger to which a hard steel ball was added to the tip as shown, and an iron frame with a steel plunger. In both cases, the bell was audibly sounded. While the sound was weak for the first configuration, it was much louder for the second.Theoretical considerations suggest that significant forces (on the order of several newtons) should be achievable with a solenoid provided with a coil such as that of Fig.

9 (left) and a plunger and frame made from a material having a much higher permeability than is now commercially available for 3-D printing.Table 1: Solenoid pull force test results.Plunger MaterialFrame MaterialPull Force (gf)SteelFe‐filled PLA 18Iron‐filled PLAFe‐filled PLA1SteelIron 32Iron‐filled PLAIron1 gConclusionsWe are developing Fiber Encapsulation Additive Manufacturing, a novel multi-materialAM process. Though FEAM can be used in principle with a variety of materials, our current focus is thermoplastic polymer as the matrix and metal wire as the encapsulated fiber, with the goal of enabling the direct, monolithic fabrication of integrated, active electromechanical and electronic devices. For such devices, we are also developing an electrically conductive polymer composite to form junctions between wires.

For such devices, we are also developing an electrically conductive polymer composite to form junctions between wires. Though much work remains to be done, we have demonstrated good progress toward overcoming the three main challenges of FEAM: achieving good centering of the wire within the extruded polymer with small radii and reasonable speeds; the ability to controllably start and stop the wire to produce short segments; and intra- and inter- layer junctions between wires with low resistance. We have also demonstrated a number of simple devices ranging from solenoids and voice coil actuators to capacitive force sensors and LVDTs.The primary application driving our work is the fabrication of soft robotic componentswith built-in actuators (e.g., electromagnetic, which are widely-used and can have excellent performance), sensors (inductive and capacitive), and interconnects.

However, other potential applications include rapid, freeform, fabrication of printed circuit boards, wearable electronics [Stoppa, 2014], stretchable electronics, defense systems such as UAVs with built-in antennasCurrent (A)0.480.480.510.46[Wang, 2015] and wing de-icing, and haptic displays. Our near-term research plans include improving the performance of junctions, building coils (such as for solenoids) in a layered fashion, and with integrated high-permeability magnetic materials, creating wire-elastomer structures for soft robotic and stretchable applications, and optimizing wire feeding and cutting.AcknowledgementsThis paper is based on work supported by the National Science Foundation under grant no. 1317961. Research was performed at the Laboratory for Additive Manufacturing, Robotics, and Automation, Department of Mechanical Engineering, Southern Methodist University, Dallas, Texas.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperADAPTING A DESIGN FOR ADDITIVE MANUFACTURING WORKFLOW TO ACCOUNT FOR CONTINUOUS CARBON FIBER REINFORCED PARTSGavin Adams Additive Manufacturing and Design The Pennsylvania State University University Park, PANicholas Meisel Engineering Design and Innovation The Pennsylvania State University University Park, PAAbstractThe use of continuous carbon fiber (CCF) reinforcement in material extrusion 3D printing has the potential to revolutionize the material extrusion field of additive manufacturing. Notably, the Markforged X7 system utilizes this CCF reinforcement with the aim to produce parts with mechanical results rivaling or surpassing those of aluminum.

Notably, the Markforged X7 system utilizes this CCF reinforcement with the aim to produce parts with mechanical results rivaling or surpassing those of aluminum. However, due to certain constraints with the deposition of CCF in material extrusion parts, such as an inability for CCF to be deposited throughout layers in the Z-direction, traditional design for additive manufacturing (DfAM) techniques need to be reevaluated. This paper will explore (1) how existing DfAM considerations (e.g., topology optimization, functional integration, minimum feature size, etc.) can be tailored to CCF and (2) how an existing DfAM workflow can be adapted to account for manufacturing limitations specific to the deposition of CCF.

The research is demonstrated through a hoist sling case study, which highlights the importance of considering fiber orientation and routing in the design stage to ensure accurate CCF reinforcement and achieve ideal mechanical results relative to the loads associated with the part. The result is an initial, potentially valuable workflow for designing CCF parts to be created using AM.1. IntroductionFor years, industries such as aerospace, automotive, and medical have sought to leverage the various advantages additive manufacturing (AM) has to offer while also searching for strong, lightweight structures. One of AM’s main advantages is its ability to create functional components with complex geometries that are difficult to manufacture by conventional and subtractive methods, thus shortening the design-manufacturing cycle and reducing production cost [1]. Material extrusion is the most widely used AM technique due to its low cost, minimal waste, and consistent accuracy [2].

Material extrusion is the most widely used AM technique due to its low cost, minimal waste, and consistent accuracy [2]. Material extrusion also requires no chemical post processing, less capital investment, and is generally considered to be a cost-effective process [3]. The material extrusion printing process has primarily served as a prototyping tool for designers; designers have, for the most part, stayed away from creating functional production parts. One of the main reasons for this is that the resultant parts tend to exhibit poor tensile and thermal performance [4]. However, recent advancements in multi-material extrusion have led to the development of processes capable of producing robust fiber-reinforced parts. Fiber reinforcements have shown to significantly increase the strength and stiffness of a part and produce components with mechanical properties similar to and sometimes exceeding aluminum [5].

Fiber reinforcements have shown to significantly increase the strength and stiffness of a part and produce components with mechanical properties similar to and sometimes exceeding aluminum [5]. This technology is offered commercially in the Markforged X7 printer, where one nozzle expels a chopped carbon fiber and nylon matrix, called Onyx, while a second nozzle deposits reinforcing continuous fibers [6–8]. The two nozzles do not extrude material simultaneously; one stops while the other is extruding. Fabricating a fiber layer starts when the bare end of the fiber is laid down and ironed in with the flat tip end of the print630head. This ironing action of the print head changes the fiber filament from a circular to an elliptical cross section geometry after being deposited onto the print bed [9].

This ironing action of the print head changes the fiber filament from a circular to an elliptical cross section geometry after being deposited onto the print bed [9]. When the routing of continuous fiber for the layer is almost complete, there is a blade system built into the extruder that cuts the strand before depositing that remaining length onto the build surface. The adhesion force of the continuous strands of fiber on the part pulls the remaining fibers through the nozzle until it is fully passed through. It is also worth noting that the start point of the fiber layer is moved for each layer so there is no corner which would serve as a weak point in the structure [9].To achieve favorable mechanical results in CCF material extrusion, the internal fibers must be laid in an orientation suitable for the applied loading conditions.

A more definitive design for additive manufacturing (DfAM) workflow needs to be developed describing how to produce an optimal design alongside case studies showing how to follow this process. To develop this, a material extrusion workflow should be established and tailored towards the new design challenges faced by CCF. S. Yang notes a series of design methodologies tailored towards AM and proposes a workflow through which successful products can be manufactured [10]. This framework is broken up into four sections: Design Specifications, Design Process, Process Constraints, and Redesigned Structure. This general design framework is especially interesting and relevant towards CCF as it initializes design from the perspective of functionality. Many DfAM workflows, by contrast, focus on optimizing an existing model.

Many DfAM workflows, by contrast, focus on optimizing an existing model. Design optimization is usually not entirely effective for CCF applications due to process limitations such as the ability to route fibers in specific orientations, limitations in available fiber lengths and material properties, and the need to align the fiber reinforcements with the loading direction to achieve ideal mechanical performance.In conjunction with design methodologies, it is important to consider the materials used in CCF printing. This knowledge can aid in achieving ideal performance from a design. By understanding the properties of these CCF-relevant materials and how they interact with the continuous fiber reinforcement process, designers can make informed decisions when creating functional end-use components. As an example, the Onyx matrix material in the Markforged X7 system exhibits a tensile strength twice as strong as ABS due to its micro-carbon reinforcement [11].

As an example, the Onyx matrix material in the Markforged X7 system exhibits a tensile strength twice as strong as ABS due to its micro-carbon reinforcement [11]. When used with continuous fiber, Onyx functions more like a matrix as its primary purpose is to bind, protect, and transfer the load to the stands of continuous fibers [1]. The reinforcing fiber options offered by the Markforged X7 include fiberglass, Kevlar, high-strength high-temperature (HSHT) fiberglass, and carbon fiber. Each fiber option has specific applications and can drastically change the characteristics of a part. Fiberglass can provide increased strength; the manufacturer claims that it exhibits 2.5x increased flexural strength and 8x stiffer properties compared to onyx in dog bone specimens. Kevlar provides abrasion resistance and is utilized in parts that experience repeated and sudden loading. This material is as stiff as fiberglass yet is more ductile.

Kevlar provides abrasion resistance and is utilized in parts that experience repeated and sudden loading. This material is as stiff as fiberglass yet is more ductile. HSHT Fiberglass produces parts with the strength properties of aluminum and has an increased heat tolerance [12]. The manufacturer claims that it is 5x stronger and 7x stiffer than Onyx. The final reinforcing fiber is carbon fiber, which offers the highest strength-to-weight ratio. The manufacturer claims that it is 6x stronger and 18x stiffer than Onyx [12]. Two different types of reinforcing fiber infills can be used with each reinforcing fiber type; the two fiber infills available within Markforged’s slicing software, Eiger, are isotropic and concentric fill. Figure 1 is a cross- sectional visualization of the two different fill patterns.631Conventional material extrusion printers offer a similar fill pattern to isotropic fiber fill.

Figure 1 is a cross- sectional visualization of the two different fill patterns.631Conventional material extrusion printers offer a similar fill pattern to isotropic fiber fill. The fibers are routed back and forth in a zig-zag pattern to simulate the unidirectional layers of a laminated composite. The fiber layers are rotated by 45 degrees to achieve unidirectional strength within a fiber group, though this 45-degree angle can be changed if needed. This option also traces concentric rings around all walls to improve wall strength. By contrast, concentric fiber fill lays the fibers only around the perimeter of a wall and has an Onyx matrix. This helps to resist bending about the Z-axis and strengthens the walls against deformation. The number of fiber shells can be specified by changing the number of concentric fiber rings within the software [13,14]. Continuous carbon fibers are composed of brittle strands that can endure loads when subjected to tension, bending, and compression.

Continuous carbon fibers are composed of brittle strands that can endure loads when subjected to tension, bending, and compression. However, they demonstrate ideal performance when utilized under tension. To withstand tension, fibers should be running along the length of the part and considered “stretched” when under force. To withstand a bending force, a sandwich panel should be utilized to overcome the outside face being in tension while the inside face is under compression. Compression loads need to have their force distribution considered. Fibers should serve as a scaffold under the load, thus distributing the load along the fiber’s path [13,14]. When defining the tensile properties of CCF, the fiber volume fraction plays a critical role in determining the part’s mechanical properties. The fiber volume fraction is the ratio between the volume of continuous fiber strands and the volume of the entire part.

The fiber volume fraction is the ratio between the volume of continuous fiber strands and the volume of the entire part. While there are other considerations to determining a part’s true mechanical performance, such as fiber fill pattern, these values provide an estimate as to how a part will perform in tension. K. Saeed attempted to characterize the material properties of continuous carbon fiber. In his studies, he performed a series of tensile tests on dog bone specimens with CCF strands and altered the part’s fiber fraction volume [15,16]. This study’s findings found a linear relationship between the fiber volume fraction and tensile strength of the part. The findings were also similar for elastic modulus. At a fiber volume fraction of 20%, CCF parts exhibit tensile properties like aluminum [15].Figure 1. Fiber Fill PatternsConsidering this body of existing research, there is still a need for a DfAM workflow which integrates the use of CCF.

Fiber Fill PatternsConsidering this body of existing research, there is still a need for a DfAM workflow which integrates the use of CCF. To address this need, this paper adapts existing best practices in DfAM to better support the creation of CCF products. Each step of this methodology will be discussed and the key considerations for each stage will be defined. The methodology will be demonstrated through a relevant case study, specifically the redesign of a load-bearing component on a hoist sling. Finally, limitations within the case study application, future research areas, and their potential impact on the workflow will also be discussed within the conclusion.6322. Adapting a DfAM Workflow for CCFThe design workflow, shown in Figure 2, has been developed by adapting the methodology proposed by S. Yang for AM-enabled design. This workflow is specifically tailored towards CCF and aims to provide a structured process for designing a successful solution for CCF reinforcement.

Yang for AM-enabled design. This workflow is specifically tailored towards CCF and aims to provide a structured process for designing a successful solution for CCF reinforcement. This adapted workflow offers a systematic approach for creating CCF products by considering the unique characteristics and challenges of CCF materials and the associated manufacturing process.Figure 2. DfAM Workflow for CCFUtilizing this design workflow can serve as a tool to help engineers the enhanced mechanical properties CCF can enable in end-use products. Following this workflow can also help reduce design error by characterizing and understanding the effects printing parameters and design decisions have on the final properties of the 3D printed composite parts for specific applications [8]. This workflow is split up into two different sections: standard material extrusion and CCF633reinforcement.

This workflow is split up into two different sections: standard material extrusion and CCF633reinforcement. The primary focus for the two categories is to clarify each process’s problem definition while accounting for its respective process constraints. The material extrusion section of the workflow is designed to adhere to the constraints of printing a part solely out of Onyx, while the CCF section aims to develop a multi-material solution that exploits the multi-material capabilities of the Markforged X7. The goal of splitting up the workflow is to ensure each specific design solution successfully captures the intended problem without introducing bias of differing process constraints. Intrinsic design decisions may be missed if attempting to capture both processes in one iterative workflow.2.1. Design Specification Clear specifications are important to engineering solutions.

Intrinsic design decisions may be missed if attempting to capture both processes in one iterative workflow.2.1. Design Specification Clear specifications are important to engineering solutions. A typical performance requirement for designs requiring the use of the Markforged X7 may include (1) a load capacity, (2) safety factor, and/or (3) number of duty cycles. Clearly understanding the functional requirements is also critical to ensure successful performance of the design and may include mating features, final part weight/size, or user features (e.g., an ergonomic way to handle a part).2.2. Problem Definition Developing a clear problem definition is a vital step to this DfAM workflow. It establishes the foundation for design thinking and influences critical design decision-making. This initial problem statement should be outlined without the influence of design constraints for CCF and be strictly focused on designing a solution for a specified material extrusion printer.

This initial problem statement should be outlined without the influence of design constraints for CCF and be strictly focused on designing a solution for a specified material extrusion printer. The problem definition should dictate a clear understanding of the specific issue the design will solve. This section defines the problem to be solved by outlining the design challenges, incorporating functional requirements and design constraints, and setting the stage for the development of a solution.2.3. Process Constraints (Material Extrusion) This section captures the constraints of the printing process. These constraints must be balanced with the design process workflow, which is why they are shown in parallel. These process constraints include material properties, printing speed, build volume, minimum feature size, and build orientation.

These process constraints include material properties, printing speed, build volume, minimum feature size, and build orientation. For both the material extrusion and CCF Process Constraints sections of this DfAM workflow, the chosen system is assumed to have a print volume of 13.00”x10.63”x7.87” as it is the volume of the Markforged X7. The minimum feature size constraint is a design principle that can be carried across most material extrusion printers. Some general design principles for minimum feature size includes a minimum wall thickness ranging between 0.047” to 0.06” and hole diameters should be greater than 0.04” in order to retain a circular shape [17].2.4. Design Process (Material Extrusion) Function Integration. Functional solutions, such as part consolidation, may be applied at this step. A “shell” design is created initially which achieves critical dimensions [10]. Other functional decisions are integrated into the design based off the problem definition.

A “shell” design is created initially which achieves critical dimensions [10]. Other functional decisions are integrated into the design based off the problem definition. The hierarchy of design decisions are a qualitative choice for the designer and should be weighed based off the defined problem. Other features such as threads and inserts are included based off the workflow. These may not yet be required after the initial pass, but a design solution check could require the engineer to general a new mating solution.634Structure Optimization & Design Practices. After a functionally successful design that adheres to mating conditions and other specifications is completed, structure optimization methods should be applied to achieve enhanced performance such as lighter weight, better heat dissipation, or improved dynamic properties [10]. As with function integration, the structure optimization hierarchy is weighed as a function of the initial problem definition.

As with function integration, the structure optimization hierarchy is weighed as a function of the initial problem definition. Topology optimization, a form of structure optimization, should be considered during this step. Leveraging the functionally successful design to produce a design which is lighter weight is advantageous and an effective step as AM provides near unlimited geometric complexity to a design. A parametric design should then be produced from the topology optimization results as an easily manipulable computer aided design (CAD) model is desirable in later stages of this workflow. This can be done by comparing the topology optimization results with the existing CAD model and removing material until a relatively similar representation is met. Though this may reduce the optimality of the final design, it is often still preferable to the initial designed topology.

Though this may reduce the optimality of the final design, it is often still preferable to the initial designed topology. It may also require some initial effort to set up a parametric design, but it can ultimately save time and effort in later stages of the workflow. As an example, a topology optimization mesh may contain features which are under the minimum feature size threshold and do not adhere to the process constraints of the material extrusion printer. Manipulating the mesh to allow for proper sized features is challenging and unwieldy; more rapid design modifications can be achieved with a similar parametric model.Design Solutions Check/Design Validation. Design solution check involves an analysis of the product to ensure that it meets the functional requirements and design constraints determined earlier. It also verifies that the design solution is feasible and will produce the desired results.

It also verifies that the design solution is feasible and will produce the desired results. The approach for a design solution check may vary depending on the specific application, especially for additive manufacturing. Finite element analysis (FEA) within most CAD software is a challenging step due to the AM material not functioning as a uniform homogenous solid. Designers must determine the necessary assumptions and acceptable format for defining a successful solution. This is an iterative step and may take a several attempts. A feedback loop to function integration or structure optimization may be necessary. If topology optimization was utilized in the Structure Optimization phase, FEA should still be conducted on the final redesigned structure to determine stress concentrations and identify how forces are dispersed within the part.

This step also serves a formal check to review the CAD model, confirming the part meets the process parameters and design specifications successfully.2.5. Enhanced Problem Definition After the completion of the previous Design Solutions Check/Design Validation step, the DfAM workflow now transitions to the CCF design process. The main objective of the CCF section is to integrate the Markforged X7’s multi-material capabilities and enable the use of reinforcing fibers. This section differs from the previous material extrusion section because it requires a more complex design process that leverages the advantages of CCF AM. This area of the design process requires an in-depth understanding of the material properties, print parameters, and design principles to create an optimal design that meets the enhanced problem definition. This section is also critical to ensuring the final design meets all functional requirements and is feasible for the CCF AM process.

This section is also critical to ensuring the final design meets all functional requirements and is feasible for the CCF AM process. The enhanced problem definition for CCF should be defined at this step; problem definition for CCF will include the combination of a lightweight, robust design with the enhanced mechanical properties that the process offers. The material extrusion workflow should have set the stage for how a solution will need to be enhanced with CCF. For example, the material635extrusion solution may need CCF to enable the mechanical performance in the design specification. CCF may also be utilized simply to provide additional confidence in a design due to the increased durability of the material.2.6. Process Constraints (CCF) While the CCF process constraints share similarities with material extrusion printing, it is important to note that they differ in terms of the ideal build orientation. The printer is limited to the depositing in the XY plane for CCF.

The printer is limited to the depositing in the XY plane for CCF. Also, CCF slows down the printing process thus increasing build time. Designers must find a balance between incorporating the right amount of CCF within the part to address this balance. Some layers may require increased isotropic CCF while others may only need concentric rings or no fiber at all. The selection between concentric and isotropic fiber fills is dictated by the type of force applied to the part. Concentric fiber rings carry the load path in a specific direction and work best when routed in tension with the load direction. Isotropic fibers provide best reinforcement in all directions resulting in a uniform distribution of the stress throughout the part.

Isotropic fibers provide best reinforcement in all directions resulting in a uniform distribution of the stress throughout the part. This makes isotropic fibers ideal for uniform stress distribution and balanced parts that need strength in all XY directions [12].General minimum feature sizes still follow the material extrusion design principles, but if a region requires CCF, the minimum feature sizes are increased. The minimum fiber reinforcement width for an open feature should be 0.15”, and 0.11” for a closed loop. Also, the minimum part height should be at least 0.04” to produce a part at least nine layers thick. This is due to the slicing software requiring four roof and four floor layers of Onyx above and below fiber groups. The smallest area to reinforce would be a square region of 0.14 inches² and a minimum post diameter of 0.38”. Sometimes, holes are too small to reinforce due to the given number of concentric fiber rings and minimum fiber length.

Sometimes, holes are too small to reinforce due to the given number of concentric fiber rings and minimum fiber length. After slicing and seeing a hole missing concentric fiber rings, this constraint may be the reason fiber is missing. Increasing the concentric fiber rings in Eiger can address this issue [13,14]. While press-fits, close-fits, and free-fits can be utilized in both CCF and material extrusion printing, they can prove to be particularly advantageous for joining split CCF parts. For material extrusion printing, a key DfAM objective is to minimize part count. However, in CCF DfAM of functional artifacts, the emphasis shifts to splitting parts in order to arrange fibers effectively. As a result, these fits play a crucial role in joining segmented parts together. Nominal clearances for the designed CAD model should be 0.000”-0.002” inch for press-fit, 0.002”-0.004” for close-fit, and 0.004”-0.008” inch for free-fit.

Nominal clearances for the designed CAD model should be 0.000”-0.002” inch for press-fit, 0.002”-0.004” for close-fit, and 0.004”-0.008” inch for free-fit. Press fitting will require some force applied via cold pressing to assemble, close fitting can be assembled or disassembled by hand with minimal clearance, and free fits will allow for parts to slide and/or rotate easily when assembled [13].2.7. Design Process (CCF) Function Integration. Function integration may not be needed on the first iteration, but it will be necessary and serve as a check before moving to special reinforcement strategies on future iterations. Most function integration accommodations may have been applied already in the material extrusion cycle; however, feedback loop design modifications for CCF may eliminate this integration and need to be readdressed. New areas of function integration may arise based off this feedback loop as well.Utilize Special Reinforcement Strategies.

New areas of function integration may arise based off this feedback loop as well.Utilize Special Reinforcement Strategies. Special reinforcement strategies for CCF should initially focus on creating a design where the load is directly transferred through the fibers. To achieve this,636the previous solution may need to be split into multiple components or printed in a different build orientation. Some examples of reinforcement strategies include directing the fibers with ribs, reinforcing the z-axis strength with other parts such as a fastener or sandwiching structures. This step may serve as a step backwards from standard material extrusion. While topology optimization is a worthwhile step for material extrusion, concessions in the original optimized design may need to be made to account for special reinforcement strategies needed for multi-material CCF.Define Fiber Reinforcement. With a clear path to direct the fibers defined, the reinforcement strategy should begin.

With a clear path to direct the fibers defined, the reinforcement strategy should begin. There are two main considerations for fiber reinforcement: direction layers and direction type. The designer should consider which regions require fibers. Some sections of a design may not carry a load or may be able to produce a successful design using only the single Onyx material. Other areas may require CCF at every layer; this is where the fiber fill should be considered, balancing between concentric outer rings and fully isotropic layers. Choosing between fiber fill paths can be done by analyzing the type of load for each component.Design Solutions Check/Design Validation. Defining an acceptable means by which to identify a successful solution is important due to the non-homogenous nature of CCF additive manufacturing.

Defining an acceptable means by which to identify a successful solution is important due to the non-homogenous nature of CCF additive manufacturing. The CCF DfAM approach builds upon the design principles and best practices established in the material extrusion section, but also introduces new challenges and considerations specific to the use of CCF. Multiple iterations through the design workflow are recommended to ensure all possible orientations, fiber directions, and part splits have been considered.2.8. Problem Solution The output may be introduced in the form of a CAD model, STL, model-based definition (MBD), or 2D drawing format. The MBD or 2D drawing can serve as a tool to display functional requirements, inspection, or capture any post-processing requirements. It Is worth noting that rarely is an engineering solution complete after the first pass. This workflow may need to be repeated after the first print of the design.

It Is worth noting that rarely is an engineering solution complete after the first pass. This workflow may need to be repeated after the first print of the design. This solution may be benchmarked against the material extrusion solution or previous revisions of the design. It can prove to be a worthwhile practice to step away from a design and compare solutions to help ensure nothing obvious was missed.3. Case StudyA relevant case study was selected to demonstrate the use and effectiveness of the proposed framework for designing CCF parts. It will show how the framework can be used to identify key design parameters, such as defining fiber orientation and part splitting, and how to utilize these parameters effectively to improve the overall performance of a final part. By using an industry- focused application, this case study provides insight into the practical applications of this framework and how it can be used to produce high performance CCF parts.

By using an industry- focused application, this case study provides insight into the practical applications of this framework and how it can be used to produce high performance CCF parts. This case study will be focused on applying the proposed DfAM workflow towards the design of a load-bearing component of a hoist sling. This hoist sling will be used to lift a machined housing around an industrial shop floor between different machining operations. Safe lifting of this component is critical as not only is it an expensive and long-lead component, but it also weighs upwards of 125 pounds. Failure of any part on the hoist sling will likely damage the product and is a safety risk for people and machinery. Figure 3 shows the existing weldment through which the lifting component must successfully mate, alongside a representation of the housing to be lifted with the637weldment assembly. The interface lifting surface is highlighted in the model.

The interface lifting surface is highlighted in the model. For the additively manufactured lifting component to successfully mate with the interfaces on the weldment assembly, it must align with two ball-lock pins and fit withing an opening between two plates.Figure 3. Weldment Assembly and Lifting Surface3.1. Design Specification The requirements for a redesigned hoist sling are as follows:A redesigned sling should be able to support a load of 125 pounds. This accounts for 33% of the working weight with a desired safety factor of 2.5.Due to the nature of the industrial environment this is used in, the design should have maximum impact resistance due to it being subjected to impacts or rough handling.

Measures should be taken to enhance the toughness of the final part by material selection and minimizing regions susceptible to breaking.The AM solution should have mounting points compatible and consistent with the hoist sling weldment and not drive a redesign of this subassembly. The part should reliably mate with these two ball-lock pins and double shear interface.The AM solution should interface on the defined contact surface in Figure 4. • A redesigned component should not damage the lifted component when in use or during assembly and disassembly.3.2. Problem Definition The material extrusion solution will be examined against the performance and functional requirements. This initial approach will attempt to create a part that mates to the critical surfaces yet is able to support the required load of 125 lbs., as determined by an initial FEA estimate. The structure should also be optimized for a material extrusion printing process.3.3.

The structure should also be optimized for a material extrusion printing process.3.3. Process Constraints (Material Extrusion) The main process constraints associated with this solution are print bed size and build direction. The correct build orientation is critical to both printing this design in one part and providing ideal part strength. The decision between these two trade-offs should be based off the design specification. For this application, a part with advantageous mechanical performance is preferred over reducing support material. The part requires two holes with a diameter of 0.500” to join the638part to the weldment assembly at the double shear interfaces. This mating requirement will not create an issue with minimum feature size. This design should have a free-fit interface with the two ball-lock pins to ensure easy installation and removal of the pin and consistent mating with the weldment assembly due to tolerance stack-ups.

Meanwhile, maximum print volume will drive the allowable surface contact between the hoist sling component and machined housing. This design solution should attempt to produce as much surface contact as possible with the machined component to help disperse the load applied while lifting the machined housing.3.4. Design Process (Material Extrusion) Function Integration. Figure 4 shows the initial concept which applies the function iteration of the design workflow as it mates with the weldment assembly and interfaces with the machined housing. For this design there are two main interface regions: the pin interface and the mounting surface. For the pin interface, this design provides two holes which align with the holes of the weldment assembly while also providing a slip-fit interface with the ball lock pins. The pin interface thickness is driven by the opening in the weldment assembly and was selected to have a nominal 0.010” clearance between the two mounting faces.

The pin interface thickness is driven by the opening in the weldment assembly and was selected to have a nominal 0.010” clearance between the two mounting faces. This will allow the bracket to be easily inserted into the groove of the weldment assembly while also having minimal movement when in use. The part then transitions towards the mounting surface. A shelf was designed for the housing assembly to rest on. There is a lip between the mounting surface and the transition region to help keep the machined housing captive during lifting. A 0.040” clearance from the outer diameter of the machined housing was used to account for the tolerance stack-up of the weldment assembly while also providing an easy and consistent fit with the machined housing. Maximum surface area on the mounting surface was designed to help distribute the load when the housing is mounted; as previously discussed, this surface area was dictated by the maximum build volume of the printer.

Additional mounting surface also helps the machined housing rest securely on the designed part. No minimum feature sizes are of concern with this design as all holes and walls are well over the printer’s minimum feature size constraints.Figure 4. Initial Material Extrusion AM SolutionStructure Optimization & Design Practices. The structure optimization phase utilized topology optimization and performed through Autodesk Fusion 360. For this topology optimization design, the objective was to minimize mass subject to the loading conditions and factor of safety discussed in Section 3.1. Figure 5 displays the topology optimization mesh results with the initial concept transparent, alongside it is the parametric redesigned structure based off these results. The intent with the topology optimization was to determine areas to remove unnecessary material and utilize639the free complexity of AM.

The intent with the topology optimization was to determine areas to remove unnecessary material and utilize639the free complexity of AM. This new structure provides a solution which not only has less material, but also will print faster.Figure 5. Topology Optimization Results and Redesigned Structure with FEA (Regions of Higher von Mises Stress Noted in Green)Design Solutions Check/Design Validation. The iterative feedback loop between the Structure Optimization and the Design Solutions Check helped determine the ideal build direction. While FEA within Solidworks is not considered reliable at predicting design AM parts tensile performance due to parts being anisotropic as opposed to isotropic [18], it is a valuable tool to determine how the force will be transmitted through a component with an applied load. Utilizing this to visualize the areas of stress concentration are critical to determining the proper build direction to mitigate delamination between layers.

Utilizing this to visualize the areas of stress concentration are critical to determining the proper build direction to mitigate delamination between layers. The rightmost image in Figure 5 displays the stress-strain FEA results which shows the region bearing the largest stress is at the pin interface. This was deemed the region which should provide the most strength. Because of this, the build direction was selected to be perpendicular to the pin interface surface. This will ensure that the region which experiences the largest load will not be susceptible to delamination between those layers. The design tradeoff with this selection is there will be an increase in support material due to the overhanging features on the mounting surface. To minimize the support material, 45-degree extrusions were created in the transition region as this is the minimum angle which does not drive support generation in Eiger.

To minimize the support material, 45-degree extrusions were created in the transition region as this is the minimum angle which does not drive support generation in Eiger. The design specification to create a part for maximum part strength over reducing support material guided this design decision. The region between the mounting surface and transition region also experiences some load and will be susceptible to delamination in this region. To increase the part’s strength in this region and minimize the possibility of delamination, the fillet size was maximized to provide as much material as possible. Figure 6 displays the build surface for the material extrusion design alongside the required support material and a visualization of the impact of the 45-degree angled extrusions.Figure 6. Material Extrusion Solution Loaded onto Build Plate in Eiger6403.5.

Material Extrusion Solution Loaded onto Build Plate in Eiger6403.5. Enhanced Problem Definition After analyzing the material extrusion design, the decision was made to create a solution that incorporates CCF to further improve the mechanical properties of the design and provide increased robustness and reliability. As detailed previously, CCF offers improved strength, stiffness, and toughness compared to an Onyx-only solution. This CCF-tailored solution should use best practices in both fiber layering and direction to produce a robust and reliable design. Following the proposed CCF design workflow, reinforcing fibers should be laid in directions that enable it to achieve the maximum mechanical properties of the material and printing process.3.6. Process Constraints (CCF) The build direction plays a critical role in the process constraints for CCF, as reinforcing fibers must be deposited in the XY plane.

Process Constraints (CCF) The build direction plays a critical role in the process constraints for CCF, as reinforcing fibers must be deposited in the XY plane. Because of this, fiber orientation is key and drives most decisions during the design process. To correctly select and design a part with CCF, the force directions and types of forces in each region must be correctly defined. For this part, the pin interface region will be experiencing a load under tension while the mounting surface will be experiencing a uniform bending force. The pin interface should have concentric fiber rings routed to support the load while the mounting interface should have isotropic fiber fill. Because of these differing load paths, this indicates splitting the material extrusion part into two components to be the best approach. This will allow for having two separate parts each designed with the correct fiber routings for their specific load paths.

This will allow for having two separate parts each designed with the correct fiber routings for their specific load paths. A strategy of mating these two components together will also be necessary in this design stage. Utilizing a press-fit and close-fit interface will be an efficient way to adhere the parts together and successfully transfer the lifting load on the weldment assembly to the machined housing. Using and designing this fit correctly will reduce post- processing as well as create a consistent finished part.3.7. Design Process (CCF) Function Integration. As stated in previous subsection, it was determined to be necessary to split this design into two separate components due to the load paths. The pin interface region exhibits a force under tension when pulled upwards and having a load applied at the mounting surface.

The pin interface region exhibits a force under tension when pulled upwards and having a load applied at the mounting surface. The transition region is where the force changes to bending as the mounting surface pulls downward while the lifting force of the weldment assembly pulls upward. This transition region and differing force types is what drives the requirement to no longer design this in one piece and instead create two different components with two different fiber path routings. The split structure for CCF is shown in Figure 7. Part One was designed with the approach of providing a part which delivers concentric fiber rings around the pin holes while also performing well under tension. The structure of Part One is similar to the pin interface region of the material extrusion solution, except it does not have a transition region. This ensures the entire part experiences a uniform load under tension and the part will be strongest where it experiences the largest load.

This ensures the entire part experiences a uniform load under tension and the part will be strongest where it experiences the largest load. It is also structured in a way to press-fit with the female square interface on Part Two and have press-fit dowel holes for a double shear interface to secure the parts together in the interface region. The functional parameter of the part width was kept the same as the material extrusion solution to provide a similar mounting scheme in the weldment assembly.Part Two was redesigned to provide a region advantageous for isotropic infill. Again, the structure of the mounting surface when compared to the material extrusion solution was kept relatively similar. The important distinction is this structure is now intended to be built perpendicular to its641bottom surface. As a result, the 45-degree extrusions in the transition region were no longer required since their function was only to reduce support material.

As a result, the 45-degree extrusions in the transition region were no longer required since their function was only to reduce support material. A female press-fit region is provided to interface with Part One. The functional intent with the dowel holes is to press-fit with Part Two at their interface points. Unlike Part One, where the dowel holes can have concentric rings around their interface points, Part Two cannot have these concentric rings as it will be built in a stair-stepped manner, perpendicular to the centerlines of the holes. This is where the square press-fit interface between the two parts comes into play. Due to the moment created from the loaded part, there will be an angular force at this interface surface when the tool is in-use making it unlikely for this fit to come loose. The function of the dowel holes is then to act as a secondary interface joint and provide added rigidity to the region.

The function of the dowel holes is then to act as a secondary interface joint and provide added rigidity to the region. The interaction between the two parts will be spread across the face of the square interface instead of being distributed entirely to the dowel holes. To increase the surface area of the interface between the two parts and help disperse the load across a larger surface, the interface joint on Part Two was raised higher. This justifies not providing a solution which has concentric rings in this area.Figure 7. CCF AM Solution with 2D-Sectioned View of CCF RoutingUtilize Special Reinforcement Strategies. Splitting the design into two different components allows for the parts to be reinforced with CCF along their ideal directions for each respective part. As discussed previously, Part One provides concentric fibers through all layers. Part Two provides isotropic fiber fill for all layers.

As discussed previously, Part One provides concentric fibers through all layers. Part Two provides isotropic fiber fill for all layers. Figure 7 shows the sectioned view for Part One and how the concentric fibers route. The blue lines are the strands of carbon fiber while the white lines are the Onyx shell and infill. Figure 7 also shows a detailed view of Part Two’s isotropic layers where the fibers are routed at 45-degree angles. There are still two concentric rings defined for these layers to help provide rigidity to the structure [13].Define Fiber Reinforcement. Both solutions will be printed with 100% infill and have two wall layers before the CCF begins being deposited for each layer. Part One utilizes a sandwich panel approach for placing the layers of isotropic CCF. Isotropic fibers are used for the first five layers and last five layers. Concentric fiber is then only deposited along the remaining layers along the middle of the part to create a sandwich structure of CCF.

Concentric fiber is then only deposited along the remaining layers along the middle of the part to create a sandwich structure of CCF. This approach will help reduce the print time while also providing a sufficiently strong part. Increasing the amount of deposited CCF drastically increases print time as the extruder deposits fiber strands much slower than Onyx. Also, routing isotropic fibers for every layer will provide minimal increase to mechanical properties.For this application, the distance the dowel pins were from the edge of part also required a feedback loop with the Function Integration portion of the design workflow for Part One. If the dowels are too close to the edge of part, they do not have enough clearance from the outer shell fiber strands.

If the dowels are too close to the edge of part, they do not have enough clearance from the outer shell fiber strands. As a result, Eiger will not slice the desired amount of concentric fiber rings for the dowel pin holes.642Figure 8 displays the dowel holes too close to the edge of part for Part One resulting in this issue compared to Figure 7 which shows the concentric rings properly routed. Part Two utilized isotropic infill for all available layers. This was required due to the bending forces distributed between the dowel pin interface and loading surface. No other special accommodations were required for the fiber reinforcement in this part as the dowel holes cannot be laid with concentric fiber rings.Figure 8. Incorrectly Routed Concentric Fiber RingsDesign Solutions Check/Design Validation. Markforged’s Eiger software includes simulation capabilities to predict the mechanical results of a CCF part.

Incorrectly Routed Concentric Fiber RingsDesign Solutions Check/Design Validation. Markforged’s Eiger software includes simulation capabilities to predict the mechanical results of a CCF part. After defining the printing parameters, CCF reinforcement, loads, and anchors, the simulation software provides a safety factor and expected maximum deflection. Both parts of the CCF reinforced solution were loaded into the simulation software to verify their performance. The simulation software indicated that Part Two initially had a failure point in the transition from the mounting surface towards the square interface between the two parts. Material in this region was increased and the fillet size was adjusted to help distribute the load evenly. Regions where the critical load was indicated were investigated as areas to increase the amount of material.

Regions where the critical load was indicated were investigated as areas to increase the amount of material. Both final iteration of the parts passed their simulation checks with maximum deflections of 0.01mm and 0.89mm and safety factors of 10.74 and 2.94 for Part One and Part Two, respectively. This meets our performance requirement for the design. Figure 9 displays the output simulation results from Eiger.Figure 9. Eiger Simulation Results3.8. Problem Solution Though physical testing of the case study has not yet occurred, to validate the final design resulting from this workflow, a comparison between a traditionally manufactured part (based on the initial material extrusion design) and the two-part CCF design was conducted. The traditional design was loaded into FEA in Solidworks with 6061-T6 aluminum; a factor of safety for the design was determined to be 8.58, below the factor of safety provided by Part One, but above the factor of safety in Part Two.

This shows how a part’s performance may be improved as a result of considering CCF throughout the design workflow.6434. ConclusionThis study explored the potential of using a CCF material extrusion process to design and manufacture high-strength functional parts. A DfAM framework was proposed, analyzed, and then followed through a case study to demonstrate the importance of a design focused on fiber routing and orientation. The proposed framework also identifies how and where to account for CFF concerns within the design process. The case study in this paper provides a clear demonstration of how DfAM principles can be effectively applied to produce complex, high-performance parts towards a real-life application. As shown in the case study, a CCF design can potentially exceed the mechanical results of a similar aluminum part. Part One in the CCF design exceeded the factor of safety results seen in the metal solution.

Part One in the CCF design exceeded the factor of safety results seen in the metal solution. Part Two did not as the fibers were unable to be routed in an orientation that provided tension with the applied load; however, the part was still able to pass the factor of safety requirement.There are some limitations which must be noted, and future research should be conducted to understand their impact. The fatigue life of Onyx and CCF designs must be better understood. These parts should be closely monitored for minor cracks or other defects when used in a real-life environment. Tensile testing must be done to compare the lifecycle of these parts against a respective subtractive metal solution. The impacts of these findings should be noted and the proposed DfAM workflow updated accordingly.

The impacts of these findings should be noted and the proposed DfAM workflow updated accordingly. The feedback on the fatigue life of these parts may impact the Function Integration or Utilize Special Reinforcement Strategies sections of the CCF workflow as additional design considerations may be required. Also, the Design Solutions Check sections for both material extrusion and CCF sections should be iterated further as additional software analysis packages and design verification methods become available.5. References[1] Ning, F., Cong, W., Qiu, J., Wei, J., and Wang, S., 2015, “Additive Manufacturing of Carbon Fiber Reinforced Thermoplastic Composites Using Fused Deposition Modeling,” Composites Part B: Engineering, 80, pp. 369–378.[2] Chee Kai, C., Rapid Prototyping: Principles and Applications in Manufacturing. [3] Wong, K. V., and Hernandez, A., 2012, “A Review of Additive Manufacturing,” ISRNMechanical Engineering, 2012, pp.

 Adaptive Control Architecture for Freeform FabricationJ .C. Boudreaux Advanced Technology Program National Institute of Standards and TechnologyIntroductionGlobal competition is driving manufacturers to significantly reduce the length of the product development cycle while maintaining product quality. But this approach has the potential of exposing firms to huge design and engineering revision costs, since even a modest revision may give rise to a cascade of new change requests. It has been demonstrated conclusively, and repeatedly, that a work-faster strategy has the insidious negative effect of forcing workers to search for local optima, which are usually ones in which the burden is shifted either upstream or downstream. A more plausible strategy is to work smarter: by increasing the amount of information available at each stage, and by identifying and correcting problems as promptly as possible.

A more plausible strategy is to work smarter: by increasing the amount of information available at each stage, and by identifying and correcting problems as promptly as possible. The work-smarter strategy has been the root cause of the phenomenal growth of rapid prototyping technologies and especially solid freeform fabrication (SFF), which includes such technologies as stereolithography (SLA), selective laser sintering (SLS), laminated object manufacturing (LOM), and ballistic particle manufacturing (BPM).Though differing in details, SFF technologies share a common operational methodology 191. First, a computer-aided manufacturing (CAD) solid model of the part is constructed. This model defines the part boundary, which separates the interior (material) region from the exterior region. Second, the CAD solid is converted into a model in the STL file format, which is a de facto standard originated by 3D Systems.

Second, the CAD solid is converted into a model in the STL file format, which is a de facto standard originated by 3D Systems. The STL model approximates the part boundary by a set of polygonal (in fact, triangular) facets. This representation permits efficient algorithms for the computation of planar intersections, but the number of polygons needed to approximate curved surfaces grows very rapidIy with the geometric complexity of the surface. STL models are also subject to errors caused in part by finite-precision floating-point arithmetic 12/. Third, the STL model is mapped into the machine's coordinate frame and then sliced into vertical cross sections. The size of the step between successive slices is a scalable parameter which typically varies from 100 to 500 micrometers. If the STL model has been correctly built, each slice will contain one or more polygonal regions.

If the STL model has been correctly built, each slice will contain one or more polygonal regions. The fabrication process then produces thin polyhedra with the required polygonal profiles either by solidifying liquids or powers or by gluing together solid laminations and cutting along the polygonal edges.In this working paper, an architecture for the adaptive control of solid freeform fabrication devices will be sketched. Close-loop controllers, such as proportional-integral derivative (PID) controllers, are already in wide industrial use. Adaptive controllers are based on a feedforward scheme in which the behavior of an inner close-loop controllers is modified, or tuned, in response to predictions based on measured changes in the fabrication environment 11/, if successful, would address many thorny issues of freeform/81.

Adaptive control,499fabrication, permitting the identification and use of such process parameters as the traverse speed of the laser, optimal layer thickness, and tolerance bounds on STL models. Generally, the main advantage of adaptive control for SFF is to deepen and extend the predictive validity of fabricated artifacts for downstream manufacturing processes.The Quality-in-Automation (QIA) Project: An ExampleThere is a connection between freeform fabrication and more traditional manufacturing technologies. This connection was stated very clearly by Chua Chee Kai in /9/, "Rapid prototyping systems are like a new type of numerically controlled (NC) tool. Unlike traditional NC, however, the algorithms needed to generate rapid prototype programs are simple.

Unlike traditional NC, however, the algorithms needed to generate rapid prototype programs are simple. There are no tool changes to make, no multiple setups, no complex surfaces to follow and no pockets to mill." An interesting observation, but my experience with the NIST Quality-in-Automation (QIA) project suggests that the connection between NC machine tools and freeform fabrication may be closer than Kai suggests. As formulated by Yee and Donmez /6/, the QIA system was designed to be a "quality control/assurance system that exploits deterministic manufacturing principles ... based on the premise that most errors in the manufacturing process are repeatable The compensation strategy was achieved by implementing tool-path and predictable... " modification of the machine using a combination of kinematic, geometric, and thermal models of machine-tool errors.

" modification of the machine using a combination of kinematic, geometric, and thermal models of machine-tool errors. These models were obtained by pre-process machine characterization measurements of the machine tools involved in this project, including a Monarch VNC-75 Vertical Machining Center with a GE-2000M CNC controller and a Monarch Metalist Turning Center with a GE-2000T CNC controller. The QIA system supported two operational modes: a real-time mode and a process-intermittent mode.In real-time mode, the compensation system monitored the machine tool and the metal cutting process, using sensory data obtained from thermal, vibration, force, and ultrasonic sensors, and then using model-based adaptive machining algorithms to do error correction by modifying the machine tool path as well as the speed and feed rates.

After the appropriate error correction was calculated, as discussed by Boudreaux in /6/ and /4/, the required compensations was achieved by sending correction signals to a specialized electronic device, called the Real Time Error Corrector (RTEC), "which is inserted between the position feedback device for each axis and the machine tool controller. It alters the feedback signal to cause the machine to go to a slightly different position to compensate for the predicted errors" /6/, /10/.In process-intermittent mode, the compensation system would identify and measure part errors which were caused by the machining process, but which could not be accounted for by means of the deterministic model. For example, tool deflection or tool wear often cause errors of this kind. These form errors were detected experimentally by on-machine gauging between the semifinishing and finishing cuts.

For example, tool deflection or tool wear often cause errors of this kind. These form errors were detected experimentally by on-machine gauging between the semifinishing and finishing cuts. Once detected, the errors were corrected by either by using the RTEC feedback system with on-the-fly (and temporary) modifications of the error model for the finishing cut, or by generating a modified NC part program (understood to be coded in conformity with EIA RS-274-D) for the finishing cut, as discussed by Bandy in /6/.To use part programs in an error compensation scheme some consideration must be given500to the methods of verifying and nominalizing them. Verification is the simpler of the two, since it can be accomplished by combining an elementary syntactic and semantic analysis of the part program with the creation of solid geometric model of the finished part and all of its intermediates.

This solid model may de generated by joining the volumes created by sweeping the cutting tool profile through the tool path and then subtracting the resulting solid from a solid model of the initial blank. The part program is verified if this solid is indistinguishable from the solid CAD model of the part. Nominalization is more difficult. In fact, there is no general agreement about the conditions which a part program has to satisfy in order to be accepted as nominal. The idea is that a part program is nominal if it would generate an exactly accurate part, satisfying all of the design constraints, if it were executed on a perfectly conformant machine, that is, a machine with no systematic errors. Since interpolation methods are important sources of non-repeatability, it has been claimed that a nominal part program should only contain linear tool paths, and that all non-linear (circular or parabolic) tool paths should be replaced by a suitable linear approximations.

It has also been claimed that the segmentation problem requires the off-line programming system to allow access not only to part programs, but also to process plans and CAD data. Such a data-rich framework would allow the programming system to maintain an internal representation of the nominal program so as to correctly propagate changes throughout the segments.The QIA project suggests the following provisional sketch of an adaptive controlarchitecture for solid freeform fabrication:Adaptive Control Architecture. (1) The inner, or core, controller accepts as command input any of an infinite set of syntactically well-formed control programs. Under the action of an algorithmic translator, the controller outputs to the plant a series of actuation signals, thereby causing the plant to peiform a corresponding series of motions. In close-loop control systems, the controller has access to feedback signals which record measured deviations between the commanded and actual plant output.

In close-loop control systems, the controller has access to feedback signals which record measured deviations between the commanded and actual plant output. (2) The adaptive controller has access to the command inpu,t and feedback signals ofthe inner controller, and also to environmental parameters whose measured values have been correlated with future plant peiformance. Under the action of an the adaptive controller may respond in any combination of three algorithmic translator, fundamental modes:I. feedback adaptation mode, modifying one or more feedback signals (corresponding to the QIA real-time mode); II. command adaptation mode, in which the adaptive correction is accomplished by modifying the inner controller's command input while preserving syntactic well-formity (corresponding to the QIA process-intermittent mode); III.

translator adaptation mode, in which the adaptive correction is accomplished by modifying the inner controller's translation algorithm.in which the adaptive correction is accomplished byThe third mode poses possibilities which are often discussed in the context of control systems based on neural net, fuzzy logic or genetic algorithm technology /1/, /8/. To make further progress, the notion of programmable devices will need to be considered more carefully.501A Theory of Programmable DevicesDevices are pieces of capital equipment that are used to modify the position, shape, size or material condition of manufactured parts. Examples of devices include numerically controlled robots, and also the freeform machines, programmable fixtures, fabrication devices mentioned above.

Examples of devices include numerically controlled robots, and also the freeform machines, programmable fixtures, fabrication devices mentioned above. More precisely, a device is an piece of capital equipment which (1) may be resolved into finitely many coupled components, each of which is itself a (2) has the capacity to perform a finite, but extensible, collection of industrially device, significant operations, and (3) may occupy any of a finite number of operational states. The first point says that devices may be nested within other devices, forming structures which resemble connected graphs. Since all devices have characteristic operations, industrial devices must support a concurrent, distributed processing model. That is, every device operation reduces to a series of concurrent operations by. its component devices, the relative motions of which are precisely orchestrated.

That is, every device operation reduces to a series of concurrent operations by. its component devices, the relative motions of which are precisely orchestrated. The controller converts these commands into a precisely structured sequence of signals to the actuation system. Thus, any realistic theory of devices must include an identification of the set of elementary or stereotypic motions that the device is able to produce, and also the definition of a method for combining elementary motions into more complicated motions.tool-changing devices,A device is programmable if every motion in its behavioral repertory, however complicated, may be produced as a response to a well-formed combination of input commands to the inner controller. There is a close connection between a industrial programmable device and I/O devices such as floppy disk drives or printers. The behavior of disk drives or printers is defined in terms of the programmed commands that then can be given.

The behavior of disk drives or printers is defined in terms of the programmed commands that then can be given. For example, a disk drive can be commanded to open a file or to access a location in a file and return a piece of data, but neither command can be given to a printer, which can only be commanded to perform text or graphics output operations. A programmable device may require non-programmed machine-tending actions on the part of an operator. For example, the part to be machined may need to be properly fixtured or the cutting tools mounted in the turret, but these operations are no different in principle loading paper into the printer or putting a floppy disk in the disk drive.Command expressions can be represented as a single phrase, consisting of a main verb, the symbolic representation of the operation, and an optional list of one or more modifiers /5/.

For example, if a suitable coordinate frame has been defined, then one might imagine that a command of the form:(MOVE (TO #(1.0 2.5 -1.0»)would cause a motion from an initial position to the position specified in the TO clause, provided the specified position is in the work volume of the machine tool. Without further the point-to-point specification of the motion permits an enormous range of qualification, acceptable path geometries. This range may be reduced by additional parametrization of the command. If one associates commands with actions by using a template mechanism, then the template mapping algorithm would take all of the parameters into account when generating the native language of the device controller /3/.

The semantic interpretation of a parametrized verb phrase is the range of behavior which that phrase would cause if it were carried out by any appropriately configured device.502A Computational FrameworkThe task of any programming language for automated manufacturing is to provide a uniform programming language environment for the construction of control interfaces to industrial manufacturing processes, and an integrated system of software tools for translating product design and process planning specifications into equipment-level control programs. To provide support for the information requirement of programmable devices, an experimental Lisp based programming language, called AMPLE /5/. This system was built in the mid 1980s for use in the NIST Automated Manufacturing Research Facility (AMRF) and was also used to design the QIA Quality Controller described above /4/, /10/.AMPLE is an interpreted language.

The behavior of the interpreter may be described as an infinite loop: (1) read an expression from an input stream, (2) evaluate the expression in the context of the symbolic environment (an ordered list of symbol/value pairs), and then (3) write the expression resulting from the evaluation to an output stream. The domain of AMPLE expressions is very simple. There are atomic expressions, including characters, strings, integers, floating point numbers, complex numbers, and arrays. When the interpreter is given any of these expressions to evaluate, it returns the expression itself. Symbols are atomic expressions which has a more complicated role. When the interpreter is given a symbol, the evaluation procedure tries to match the symbol with the all of the entries in the symbolic environment. If a match is found, then value corresponding to the symbol is returned. If there is no match found, then the interpreter returns a message to that effect.

If a match is found, then value corresponding to the symbol is returned. If there is no match found, then the interpreter returns a message to that effect. In addition to atoms, the only other class of expressions are lists. When the interpreter is given a list to evaluate, it assumes that the expression is the application of a function, which must be the first component of the list, to the interpreted values of the arguments, which must be the remaining components of the list. The value obtained is then returned as the value of the original list.The value of an expression depends upon the operation of the interpreter and the current content of the symbolic environment. When initiated, AMPLE comes supplied with a set of primitive functions and variable symbols. This primitive environment may be dynamically altered by subsequent operations of the interpreter. As explained in /5/, some functions, when evaluated, cause new symbol/value pairs to added to the environment.

As explained in /5/, some functions, when evaluated, cause new symbol/value pairs to added to the environment. For example, the evaluation of an assignment expression can cause the symbol pi to be added with the associated value 3.14159, or the evaluation of a function definition expression can cause a new function to be added. All symbolic environments are persistent objects which can be stored in external files and then loaded back when needed. The normal effect of reloading an environment is to restore the prior configuration, which means that AMPLE programming is focused on the construction of snap-on workspaces. Many workspaces have been developed, three of which are directly relevant to the topic of this paper: concurrency, code generation, and real-time processing.Concurrency. There are two different issues that an industrial programming system must resolve: what operation, or operation sequence, needs to be performed, and when must the operations be initiated and terminated.

The program needs to specify both the operations and the order in which the operations need to be done. Since a single device may actually contain503several controllers, and thus different operations can be executed concurrently, an efficient representation of these programs is by means of directed graphs (digraphs).A set of basic functions for concurrency has been defined by Hoare /7/: spawn causes a new process to be created and added to the list of active processes, block causes a process to be unavailable for further processing by removing it from the active list, activate restores a blocked process to the active list, and alt, when given a set of alternatives, causes one of them to be selected. In addition, the function par causes agents to bring about explicit parallelism. When par is applied to a sequence of processes, called children, the process containing the occurrence of par being executed, called the parent, is immediately blocked.

When par is applied to a sequence of processes, called children, the process containing the occurrence of par being executed, called the parent, is immediately blocked. The children are concurrently spawned, and are then processed in the manner just defined. When all of the In many children have successfully terminated, and only then, situations it is necessary for processing agents to coordinate their behavior. This is accomplished by means of communication channels. These are special processes which are used as structural members in order to establish a communication link between two or more processing agents, some of which want to participate in an input event, and others in an output event. For example, suppose that there is a Boolean-valued symbol, say idle, in the symbolic environment which is to be used to hold the current operating condition of a NC machine. This can be implemented by communication channels.

This can be implemented by communication channels. The NC machine, actually a processing agent going proxy for the machine, is allowed to participate in output events with respect to idle, and any of a number of agents may participate in input events. The channel enforces a hand-shake discipline, that is, when the machine agent needs to output to idle, it acquires the output port, blocks all other input-requesting agents from acquiring the input port. The agent then updates idle and releases both the input port and the output port. If input-requesting agents are blocked, either by an update in process or by other inputing agents, they may, at their own option, either This example shows that queue up to peek at communication channels provide an elegant way to implement device states.the parent is reactivated.idle or simply go abouttheir business.Code Generation.

A code generator is a program which translates high-level device actions into low-level application programs in the device controller's native language. Code generation, which encourages effective re-use of existing control programs, is typically carried out in three stages. The first stage consist in the elaboration of precise definitions for the target family of devices. These definitions should group devices into classes based on their components, the operations that they can be commanded to perform, and their operating states. The second stage consists of the development of a library of program templates, each of which has been constructed by inserting parameters within carefully chosen fragments of control programs. The third, and last, stage is the definition of combinatorial rules for stringing templates together and uniformly replacing the embedded parameters with appropriate values.There are many possible template processors techniques.

One requires a distinction between a context, which is a linear sequence of characters, and algorithms or chunks of executable code, which are embedded within the context and whose interaction with it can be very complicated. The embedded code is prefixed by a selected special character, such as the at-sign @. The template processor scans each character in the source file in the normal left-to-right order and writes the character to the target file. When a special symbol is the template processor assumes that it is immediately followed by a legal encountered,504expression, which is to be passed to the interpreter for evaluation. The output to the target file In general, all is determined in the course of the evaluation of the embedded expression.

The output to the target file In general, all is determined in the course of the evaluation of the embedded expression. template processors must include a completely unambiguous algorithmic presentation of the mapping function which, given any legal template and (possibly) any legal responses from external information sources, will generate the target text in a deterministic manner.Real-Time Processing. Since all interpretive systems, need to reclaim memory, a process called garbage collection, which is a time consuming procedure, the interpreter has to have an internal real-time processor (rtp). This processor is a programmable module, acting as a (virtual) coprocessor of the AMPLE interpreter. The rtp has a limited operation set but a sophisticated set of data structuring capabilities.

This processor is a programmable module, acting as a (virtual) coprocessor of the AMPLE interpreter. The rtp has a limited operation set but a sophisticated set of data structuring capabilities. This processor needs to address two conflicting design goals: it must be able to execute at very nearly compiled language speeds, and it must have direct access to objects dynamically created by the interpreter. This conflict can be resolved by using special off-line programming services to build the objects, which are then accessed and updated using the special access functions within rtp. Moreover, to prevent garbage collection, the access and update functions must never cause the allocation of any free memory cells.Every rtp is linked to external consumer systems through ports. A port is a finite sequence of bytes which represent the remote communication between these systems together with a protocol which defines input and output behaviors of both systems.

A port is a finite sequence of bytes which represent the remote communication between these systems together with a protocol which defines input and output behaviors of both systems. A clear example is a communication system in which one system, call it the target, writes instantaneous position information and reads correction signals, and the second system reads the position signal and writes the correction signal.Any answer returned by rtp can be assessed in terms of its mathematical correctness, but real-time computing places an additional burden: the assessment of correctness also depends on a payoff function. For example, hard real-time systems are those whose payoff is a non-zero constant value if the correct value is returned within the allowed interval, and 0 everywhere else. In this case, there is a constant payoff for an answer delivered within the interval, but once the time interval has past, the answer has no value at all.

In this case, there is a constant payoff for an answer delivered within the interval, but once the time interval has past, the answer has no value at all. A different payoff function might support different strategies. For example, given a payoff function with a non-zero initial value which slowly approaches 0, then there would always be at least a small advantage to returning an answer.Discussionthe identification of basic operations, and the definition of a methodology for modifying, or adapting, these processes to specific situations. It has been proposed in this paper that the computationally difficult part of this problem can be handled by an very-high-Ievel interpretive adaptive control architecture.

It has been proposed in this paper that the computationally difficult part of this problem can be handled by an very-high-Ievel interpretive adaptive control architecture. This adaptive control family behaves as a system with a modal switch so that adaptation may be accomplished by modifications of the feedback system or the command system, both of which have been extensively investigated, or by structural modifications of the translation system of the inner controller, which is the most difficult case.The control problem for solid freeform fabrication devices has two classical aspects:505Until now, systems with this capacity have been based upon either neural nets, fuzzy logic, or the genetic algorithm. For example, a neural net consists of a network of active elements, called neurons. Each neuron may have several input connections, but all have only one output. Each input connection has an assigned weight.

Each neuron may have several input connections, but all have only one output. Each input connection has an assigned weight. The value returned by each neuron is determined by multiplying each input value by its associated weight, summing these products together, and then In this case, adaptation is a matter of by applying a sigmoid function to the computed sum. updating (that The adaptive control architecture sketched in this working draft suggests an approach to this problem through the design and implementation of what might be called an adaptation-instrumented command translator.is,incrementally revising)the associated weights.References1. Astrom, K.J. and Wittenmark, B. Adaptive Control, Addison-Wesley, 1989. 2. Bohn, J.H. "Removing Zero-Volume Parts from CAD Models for Layered Manufacturing," IEEE Computer Graphics and Applications, November 1995, 27-34. 3. Boudreaux, J.C.

2. Bohn, J.H. "Removing Zero-Volume Parts from CAD Models for Layered Manufacturing," IEEE Computer Graphics and Applications, November 1995, 27-34. 3. Boudreaux, J.C. "Code generation techniques to support device abstraction," International Conference on Manufacturing Systems and Standardization, ICMSS '91, Budapest, Hungary, 1991, 1-8. 4. Boudreaux, J.C. "A Programming Language Environment for Quality Control," International Programmable Controllers Conference, IPC'91 , 1991, 579-590. 5. Boudreaux,. J.C. in W.A. Gruver and J.C. Boudreaux (eds.), Environments for CIM, Springer-Verlag, 1993; 31-91. 6. Donmez, M.A. (ed.) Progress Report of the Quality in Automation Project for FY 90, NISTIR 4536, March 1991. 7. Hoare, C.A.R. Communicating Sequential Processes, Prentice-Hall, Englewood Cliffs, NJ, 1985. 8. Ioannou, P.A. and Sun, J. Robust Adaptive Control, Prentice-Hall, 1996. 9.

7. Hoare, C.A.R. Communicating Sequential Processes, Prentice-Hall, Englewood Cliffs, NJ, 1985. 8. Ioannou, P.A. and Sun, J. Robust Adaptive Control, Prentice-Hall, 1996. 9. Kai, Chua Chee "Three-dimensional rapid prototyping technologies and key development areas," Computing and Control Engineering Journal, August 1994; 200-206. 10. Yee, K.W., Bandy, H.T., Boudreaux, J.C., and Wilkin, N.D. "Automated Compensation of Part Errors Determined by In-Process Gauging," NISTIR 4854, 1992. 20'th506.

APTIVE LAMINATED MACHINING FOR PROTOTVPING OF DIES AND MOLDSF. A. Vouzelaud and A. Bagchi Department of Mechanical Engineering Center for Advanced Manufacturing Clemson University Clemson, SC and P. F. Sferro Ford Motor Co. Dearborn, MIABSTRACTAdaptive laminated machining is the fusion of slicing a solid model into layers and producing parts by CNC milling machines. Unlike other solid freeform fabrication processes which create the part by addition of material, adaptive laminated machining can create solid parts by selectively removing in layers. The research issues and practical limitations on shape and manufacturability are thus different from other processes. However, the biggest advantage is the ability to obtain a solid metal part such as a die or a mold directly. In this paper, the concept of this technique, and initial results and parts produced in Clemson will be presented.

In this paper, the concept of this technique, and initial results and parts produced in Clemson will be presented. In addition, future research needs and issues will be discussed.INTRODUCTIONThe main problem in manufacturing prototypes lies in the control of the delay between design and production. Until recently, a prototype of a new part required weeks or months merely to see if the shape or features would be acceptable or not. With the newer technologies developed in the past four years, a prototype can be created within hours. These new processes are grouped under the generic term free form fabrication (FFF).FFF techniques can produce parts without any shape restriction because: (i) no tooling is required and (ii) parts are built layer by layer. 3D Systems' StereoLithography, DTM's Solid Laser Sintering, Cubital's Solider, Stratasys' Fused Deposition Molding are some of the commercially available prototyping units.

3D Systems' StereoLithography, DTM's Solid Laser Sintering, Cubital's Solider, Stratasys' Fused Deposition Molding are some of the commercially available prototyping units. If texture quality and dimensional accuracy are competitive with more traditional the slices are thin, manufacturing processes like milling and turning. However, the materials used by the FFF techniques are limited to polymers (liquid and pOWder) and ceramic powders, and cannot reach the mechanical properties required for many practical uses, such as dies and molds. Therefore, these processes are limited to the production of smaller parts and those to represent shapes and sizes only.1A recent study presented at an MIT conference on Leadership for Manufacturing in 1991 (Table 1) shows that computer numerical controlled (CNC) machining matchesstereolithography in many points. The greatest advantage of FFF over CNC machining is its capacity to produce any intricate shape.

The greatest advantage of FFF over CNC machining is its capacity to produce any intricate shape. However, most, if not all, FFF processes are limited to non-metallic materials or metals with a relatively large void In addition, due to size limitations, dimensions of parts producible on these FFF units are limited fraction. to 20 inch cubed. CNC machining is limited by its programming time needed to convert a three dimension solid model to a machined part. If the programming time for CNC machining can be reduced without affecting other parameters, new applications can be developed for this conventional manufacturing process for both prototyping and limited batch production.Several other processes are being developed using paper, metal and ceramic powders and other materials as of this writing.

However, these processes are not commercially available and were not included here.291In this paper, a concept on product realization from a CAD solid model is developed by fusing FFF processes and conventional machining. This concept, called adaptive laminated machining, is based on the integration of slice generation as in FFF technologies and 21h dimension CNC machining. The overall idea is to decompose a solid model of the part into layers and then to automatically machine each layer to create the finished shape. The final aim is the automation of prototyping of dies and molds from a computer aided design (CAD) model. The approach taken to develop adaptive laminated machining and some results is the sUbject of this paper.BACKGROUNDIntroduction to FFFSeveral techniques can be used to produce prototypes, like stereolithography, solid base curing, selective laser sintering, fused deposition modeling and three dimensional printing [1-6].

All of these techniques decompose the CAD model into a number of layers, called slices. The slices generated by the slicing algorithm comprise of basic entities, such as arcs, circles and straight lines, which are used to reconstruct a solid object. Therefore the first step for all FFF processes is slicing based on the graphical representation of the part. An overall scheme presented in Figure 1 describes the process flow.The advantages in using an FFF process are: texture quality and dimensional accuracy can compete with conventional machining; no shape constraint comes from the process itself, the only limitation is associated with CAD; the manufacturing process is fully automated, and can be run untended overnight; no tooling or special fixturing is required; the production of several parts at the same time with different geometries is possible;1. 2. 3. 4. 5.

2. 3. 4. 5. and 6.the building speed is a function of both process characteristics and dimensions of the part, and it is not affected by the geometrical complexity. The main drawback of these processes is that because of the materials used to date, the mechanical properties of the parts are insufficient. Hence the applications of the processes are limited to look-alike parts rather than true prototypes with the same strength, stiffness and toughness. Secondly, the initial cost of hardware, software and training is very high when compared with conventional manufacturing techniques. These technologies utilize novel materials and concepts, and thus require the acquisition of knowledge of the materials and familiarization with the concepts (e.g., slicing) and machines (e.g., stereolithography apparatus).Table 2 compares the characteristics of different FFF processes based on the data obtained from the references.

It should be recognized that since all these processes are new, the values are based on few experimental measures, some on theoretical calculations, while those for some others are not available at this time. With time the characteristics of these processes should be better understood and evaluated.Slicing AlgorithmsSlices are 2 dimensional geometrical entities (or cross-sections) assembled sequentially to create a solid object. Since all rapid prototyping technologies fabricate objects out of slices, they are the basic geometrical entities [7]. The general procedure to produce the slices from the CAD representation of the object is shown in Figure 2. Because surface texture, and form and dimensional accuracy are important to the prototype, it is essential to keep the best CAD representation possible to minimize geometrical error. The first step in Figure 2 shows the transfer of the original CAD file into an STL format, which has become the de facto standard for FFF processes.

The first step in Figure 2 shows the transfer of the original CAD file into an STL format, which has become the de facto standard for FFF processes. This new representation approximates the surface of the object by triangles, the vertices of which belong to the original surface. The accuracy of the tessellated file can be improved by increasing the number of facets, requiring greater storage space and computation time. The slicing algorithm then utilizes this STL file and computes the intersection between the object and horizontal planes. The boundaries obtained through such an intersection, represented by polygons, is the cross-section of the object at that layer. The information on the cross-section of the object is utilized to create the part one layer at a time.292At present the concept of slicing is well established.

The information on the cross-section of the object is utilized to create the part one layer at a time.292At present the concept of slicing is well established. Many commercial CAD packages provide In addition, capabilities for slice generation from a solid model researchers at University of Dayton, Carnegie Mellon University and Clemson University [8] have been working independently to automatically convert solid models to slices. The key issue in developing such indigenous software is to provide faster computation, better control over the process and minimize errors due to slicing.to generate the software manually.ADAPTIVE LAMINATED MACHININGGiven that it is not possible to produce metal parts with FFF processes, classical techniques like machining need to be developed to reduce the time between design and production.

Although computer aided manufacturing systems (CAM) have considerably simplified process planning, they are not adapted to the complex shapes encountered in parts like dies and molds. Such tasks require the knowledge of an experienced operator. The combination of CNC machining and production of parts layer by layer can give a satisfactory answer to this specific problem.The objective of this research is to develop an integrated design for manufacturing system to prototype dies and molds using adaptive laminated machining as presented in Figure 3. Similar to slicing and cross hatching in stereolithography, tool path planning is of paramount interest in adaptive laminated machining. The core of the study will focus on issues related to tool path generation shown in the highlighted boxes in Figure 3.Slice controlled machining is based on the decomposition of the solid object into layers, and a contour describing the cross-section of the object at each layer.

A complex die or mold can thus be rapidly manufactured for prototyping purposes by breaking it up into simpler three dimensional segments (if necessary), decomposing each of these segments into layers, machining the contour for each layer, and thus creating the die layer by layer.

This process thus requires an efficient slicing process for the original solid model and a tool path generation algorithm for both rough machining as well as finish machining.The objective will be met by addressing three research issUes: (i) developing the interface between a CAD system and CNC machining through a slicing algorithm (to be developed for adaptive laminated machining) (Ii) developing an integral tool path generator to work with the slicing process and a wide variety of milling machines by using standard commands; and (iii) characterizing the quality of parts obtained by slice controlled machining to produce dies and molds whose quality will be comparable to those obtained using existing manufacturing processes (such as casting and machining).CONCEPT AND VALIDATIONBy selecting machining to fabricate dies and molds, most of the process parameters are already pre established by the working environment.

Examples in point include shape complexity, presence of cooling and/or heating circuits, fixturing constraints, strength and stiffness at critical locations in the die or mold, surface quality obtainable using CNC machining.In adaptive laminated machining, the process planning and execution of CNC machining will be completely modified. This approach leads to the creation of computer numerical controlled (CNC) machine codes for each slice generated, proceeding one slice at a time from top to bottom. This will permit the handling of any shape of a die surface with the desired texture, and dimensional and form accuracy. in conventional methods, the special features would need to be identified and generated Alternatively, individually utilizing machining rules in an expert system which is being investigated by other researchers [9, 10].The overall concept of adaptive laminated machining is presented in Figure 4. The sliced boundaries are directly obtained from the original CAD model.

The sliced boundaries are directly obtained from the original CAD model. Therefore an intermediate file like the STL format, which is a faceted representation of the geometry and used by many other·commercially viable rapid prototyping systems, can be avoided. Due to this difference with other rapid prototyping systems' approach, existing slicing algorithms are not applicable any more. However, slicing becomes easier because only the intersecting boundaries of the solid model with the slicing planes are needed, and no post processing of293data to obtain any hatching inside the contour is needed. Therefore, a built-in macro in any CAD package to intersect a plane with a solid model can be used. This implementation inside the CAD environment can benefit from flexibility and diversity of CAD commands.

This implementation inside the CAD environment can benefit from flexibility and diversity of CAD commands. Thus, the algorithms can focus on quality issues instead of trying to solve entity manipulation, intersection, and graphic representation problems.The object is drawn using AutoCAD AME extension as a solid model. Using the standard AutoCAD plane intersection algorithm the object is divided into different layers. Thus, the slicing is done within AutoCAD using its C libraries. The resultant is a slice containing the boundary of the object due to intersection of the layer with the object. The object is thus broken down into a number of layers, separated by the slice thickness specified by the planner. Each layer boundary is stored in a file, sequentially, utilizing the AutoCAD data exchange format (DXF). This file thus contains line and arc segments corresponding to each layer.

Each layer boundary is stored in a file, sequentially, utilizing the AutoCAD data exchange format (DXF). This file thus contains line and arc segments corresponding to each layer. Each line or arc segment is then defined in terms of its radius of curvature (if appropriate) and the coordinates of its end points. This data is used in a C program to generate the process plan from layer to layer and SUbsequently the numerical part program to drive the machine.The example in Figure 5 illustrates the feasibility of the method with a manually varied z increment. It is possible also to develop the slices with variable z increments to better represent the surfaces [11]. The finished part produced on a desktop 3 axis milling machine using a flat end mill is shown in Figure 6.

The finished part produced on a desktop 3 axis milling machine using a flat end mill is shown in Figure 6. The work to date clearly shows the feasibility of the process and the ability to produce a large die or mold cavity from a metal block using adaptive laminated machining.DISCUSSIONIn addition to machining limitations are expected using this approach. constraints imposed by the milling machine itself, the slicing process adds new shape restrictions. Feature like threads, grooves or convex shapes are to be machined using a more conventional technique or after refixturing. Figure 7 provides some examples of geometrical constraints in adaptive laminated machining. Once the undesired features are removed from the CAD model and the pockets are extracted, the minimum number of fixturing has to be determined as a function of the degrees of freedom of the machine tool.

If a single fixturing is desired, the solid model can be used directly by slicing program, with no further modification. Parts such as dies and molds have geometrical properties where the positive shape is extractable and these parts do not have any undercut. If the parts need refixturing, only a minimum number of fixturing needs to be used. This is an area of ongoing research and will be discussed in depth in a future paper.Several geometricalThe direction of the slicing and identification of dominant features are two other domains of research in adaptive laminated machining. The choice of the direction of slicing will relate to the minimum number of fixturing needed and thus ultimately to producibility of shapes by machining. Another issue of concern is tool path generation and tool offset necessary for different geometries.

Another issue of concern is tool path generation and tool offset necessary for different geometries. These too will be discussed in a future paper.ACKNOWLEDGEMENTSupport of this work by Ford Motor Co.through a grant to Clemson University is gratefullyacknowledged.REFERENCES1.2.3.E. Sachs, M. Cima, and J. Bredt, 'Three Dimensional Printing of Ceramic Shells And Cores For Metal Casting", Intelligent Design and Manufacturing for Prototyping, PED-Volume 50, ASME 1991. L.L. Kimble, Technology", Intelligent design and manufacturing for prototyping, PED-Volume 50, ASME 1991. E.P. Gargiulo, and D.A. Belfiore, "Photopolymer Solid Imaging Process Accuracy", Intelligent design and manufacturing for prototyping, PED-Volume 50, ASME 1991."The Selective Laser Sintering Process: Applications Of A New Manufacturing2944.5. 6.7.8.9.10.11.S.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceAN ADAPTIVE MESHING FRAMEWORK USING OCTREE DATA STRUCTURE FOR VOXEL BASED MESHESPradeep Chalavadi*Ansys Inc, 2600 Ansys Dr, Canonsburg, PA 15317AbstractWe present an adaptive meshing framework for voxel-based meshes, designed for use in various process simulations for additive manufacturing, such as thermal, distortion, grain growth, etc. The framework uses an octree data structure the meshes, and a coarsening/refinement algorithm to generate coarser and finer meshes. The algorithm preserves a 2:1 ratio of coarse to fine meshes to maintain desired accuracy. Efficient tree traversal is used for fast nodal/Gaussian solution mapping. In many cases, selective element coarsening enables the reduction of the number of nodes to be solved by the iterative matrix solver.

In many cases, selective element coarsening enables the reduction of the number of nodes to be solved by the iterative matrix solver. To maintain accuracy at boundary, the algorithm can be configured to maintain a certain level of fine mesh at boundary. When part and support mesh touch, they are automatically flagged to be not combined to be coarsened at any stage. Overall, the algorithm enables reduction of solution nodes while maintaining desired accuracy at areas of interest.to representIntroductionIn the ever-evolving landscape of modern manufacturing, Additive Manufacturing (AM) has emerged as a transformative technology, promising unprecedented flexibility, customization, and efficiency. By building intricate components layer by layer, AM processes have shattered traditional manufacturing constraints, enabling the production of complex geometries that were once thought impossible.

By building intricate components layer by layer, AM processes have shattered traditional manufacturing constraints, enabling the production of complex geometries that were once thought impossible. However, the realization of this potential relies heavily on a deep understanding of the intricate interplay between virtual design and real-world production—a challenge that has brought forth the crucial role of simulation.Simulation of additive manufacturing processes involves making various assumptions and cast the physical process into a form that can be manageable numerically and programmatically. There are multiple approaches of doing this involving varying levels of assumptions and accuracies. Particle-based simulations with μm scale phenomena delve into the intricate dynamics of powder particles during the initial stages of additive manufacturing, typically at the micrometer scale [1].

Particle-based simulations with μm scale phenomena delve into the intricate dynamics of powder particles during the initial stages of additive manufacturing, typically at the micrometer scale [1]. These simulations aim to model the behavior of individual particles, including powder flow, melting, and solidification. While offering high accuracy by capturing fine-grained powder interactions, they come at the cost of computational intensity. On the other hand, simulations that start with solidified material at the micro-level and solve each laser scan line focus on modeling the interaction between the laser and the material to predict thermal and mechanical effects accurately [2]. These methods, while faster than particle-based simulations, can still be prohibitively slow and computationally intensive to get a part level solution.

These methods, while faster than particle-based simulations, can still be prohibitively slow and computationally intensive to get a part level solution. Hatch-by-hatch coupled thermo-mechanical lumping divides the build process into discrete segments similar to1729laser hatches and considers thermal and mechanical interactions within each hatch [3]. While these methods are still very accurate they are not significantly faster than micro-level thermos- mechanical simulations. Layer-by-layer coupled thermo-mechanical lumping with total laser power flashing extends the simulation to a layer-by-layer analysis, taking into account the cumulative effect of multiple scan paths [4]. Total laser power flashing considers the heat input over an entire layer. This approach enhances accuracy by incorporating more build details while retaining computational efficiency by lumping within each layer. It is valuable for comprehensive analysis of part quality.

This approach enhances accuracy by incorporating more build details while retaining computational efficiency by lumping within each layer. It is valuable for comprehensive analysis of part quality. For a more rapid assessment, instead of doing a coupled thermo- mechanical analysis, there are approaches [5] which assume a fixed strain for a lumped layer and perform a mechanical simulation to obtain the final distortion and residual stresses at macro scale part level.Many of these approaches, especially involving part level simulations require a finite element mesh with enough resolution and quality. Since the laser melting phenomenon and the subsequent solidification phenomenon occur at a micro-scale the resolution required to capture this spatio-temporal process is very high and challenging. To solve this problem to an extent voxel based meshes provide an accurate three-dimensional grid structure that can faithfully capture the intricacies of these complex geometries.

To solve this problem to an extent voxel based meshes provide an accurate three-dimensional grid structure that can faithfully capture the intricacies of these complex geometries. Additively manufactured parts often feature intricate and complex geometries, including internal voids, overhangs, and lattices. Voxel based meshes allow for a direct representation of the digital model in a way that maintains geometric fidelity, ensuring that simulation results are as close as possible to real-world outcomes. In view of these advantages voxel-based meshes have gained popularity in AM simulations for many types of approaches discussed above. One challenge with voxel-based meshes is the amount of resolution to capture feature geometry is very high which will result in a significant computational cost. It is very hard to approach a balance of mesh resolution, solution time and accuracy.

It is very hard to approach a balance of mesh resolution, solution time and accuracy. Considering this, it is desirable to employ adaptive meshing that simplifies portions of the geometry when temperature, stress, or distortion exhibits limited variation. Combining neighboring voxels into a bigger voxel and mapping the solution variables to the new mesh element(voxel) is a known approach. In this paper we discuss an approach to achieve this in an automated fashion which is numerically efficient and gives accurate representation. Octree data structure and algorithms have been employed to implement this. The method preserves geometric boundaries to have the original fine mesh. It keeps the neighboring elements to not more than one level coarser. We also discuss the underlying assumptions and accuracy preserving features.

It keeps the neighboring elements to not more than one level coarser. We also discuss the underlying assumptions and accuracy preserving features. It also shows the necessary controls to preserve the geometry and correct for local variations.Octree data structure for mesh encodingAn octree-based data structure is a hierarchical representation commonly used formanaging spatial data and traversing three-dimensional space in an efficient manner [6]. The current method starts with a fine mesh with desired resolution for accuracy purposes. It then begins by considering the whole geometric domain with voxels as a single parent element/node of a tree structure. This parent node is divided into eight spatial elements/nodes called octants. Now, each of these further divided into another eight octants. This process is repeated until the smallest octant is equal to the original voxel size. During the sub-division process the parent child relationship is encoded into data structures.

This process is repeated until the smallest octant is equal to the original voxel size. During the sub-division process the parent child relationship is encoded into data structures. Also, when one octant encompasses a volume which doesn’t contain any voxel belonging to the original geometry it is kept from any more sub-1730division. An example of this idea is illustrated in Fig. 1 in a simplified 2D structure comprising of quadtrees instead of octrees but are very similar in nature.(a)(b)(c)Fig. 1 a) original geometric domain; b) quadtree representation of the original geometry; c) related tree representation of the quadtree divisionsThe resulting data structure will have mesh elements at leaf nodes of the tree structure.Some of the leaf nodes might, which are marked as hollow in Fig. 1 are not part of the geometry but are there for completeness of the octree structure. An example of a smaller resolution 3D mesh and corresponding octree is shown in Fig.

1 are not part of the geometry but are there for completeness of the octree structure. An example of a smaller resolution 3D mesh and corresponding octree is shown in Fig. 2.(a)(b)Fig. 2 a) Original geometric domain with color coding of octants; b) octree representation of the original geometry divisionsSince, the octree representation starts with a voxel mesh and voxel meshes are easy formanipulation when compared to unstructured meshes due to their inherent ease in combining smaller elements into larger ones within a given neighborhood, and vice versa. This versatility in element manipulation is further enhanced by the implementation of octrees, which provide a1731structured and efficient means of achieving such operations, see Fig. 3. The octree data structure serves as a key component in this context, offering a systematic approach to creating hierarchies of mesh elements.Fig.

3. The octree data structure serves as a key component in this context, offering a systematic approach to creating hierarchies of mesh elements.Fig. 3 Pattern mesh structures coarsened in a gradiated manner using octree data structure.This hierarchical organization enables effortless collapsing and expansion of meshcomponents as needed. The octree's regularized framework ensures that the mesh retains its structural integrity while undergoing refinement or coarsening, see Fig.4. Parsing the geometric information into an octree-type data structure brings about significant benefits in terms of geometric manipulation. With this structure in place, traversing through the geometry becomes remarkably convenient, allowing for precise access to specific locations that may require coarsening or refinement.

With this structure in place, traversing through the geometry becomes remarkably convenient, allowing for precise access to specific locations that may require coarsening or refinement. This combination of voxel meshes and octrees streamlines the process of mesh manipulation, facilitating the efficient adaptation of mesh elements to meet various computational needs.Fig. 4 Geometric integrity maintained during coarsening using octree data structures.1732Additive Process Simulation ConsiderationsOne of the most critical factors influencing the success of numerous additive manufacturing process simulations is the precise localization of laser power or energy. This phenomenon plays a pivotal role in determining the quality and integrity of the final result. In the context of scan-line-based simulations, it becomes evident that solution variables are particularly sensitive to the mesh arrangement surrounding the path of the laser beam.

In the context of scan-line-based simulations, it becomes evident that solution variables are particularly sensitive to the mesh arrangement surrounding the path of the laser beam. To accurately capture the thermal gradients, as well as the consequential locked-in residual stresses and distortion, it is imperative to employ a very fine mesh in the vicinity of the laser beam. This fine mesh ensures that the laser's energy deposition is faithfully represented in the simulation, thus allowing for a more realistic and predictive modeling of the additive manufacturing process. Similarly, in the case of layer-by-layer simulations, where laser power lumping is employed to model entire layers, a high-resolution mesh becomes vital around the uppermost layers. These upper layers, situated close to the rapidly solidifying molten material, require a fine mesh.

These upper layers, situated close to the rapidly solidifying molten material, require a fine mesh. This level of resolution is essential to accurately capture the thermal dynamics at play, ensuring that the simulation results exhibits the desired outputs verifiable with experimental results.However, it's not only the laser power's localized effects that demand careful consideration. The boundary surfaces of the geometry also undergo significant transformation, experiencing distortions, thermal signatures, and residual stresses. Consequently, maintaining an adequate mesh resolution along the boundary surfaces is equally vital.

Consequently, maintaining an adequate mesh resolution along the boundary surfaces is equally vital. This ensures that the impact of the additive manufacturing process is faithfully reflected in the simulation, allowing for a comprehensive assessment of the final product's quality and structural integrity.Balancing these intricate requirements, we find that there is room for mesh coarsening in areas of the geometry that lie further away from the laser beam path and the boundary surfaces. By intelligently optimizing the mesh structure using octree data structures, we can efficiently achieve the necessary mesh refinement where it matters most while economizing computational resources in regions less critical to the simulation. This work demonstrates the efficacy and speed with which octree data structures can address these multifaceted considerations. Some of these above discussed considerations are depicted in Fig. 5.Fig.

Some of these above discussed considerations are depicted in Fig. 5.Fig. 5 Additive process simulation related considerations reflected in coarsening with octree: Fine mesh elements are maintained at top layer vicinity and boundary element vicinity.1733Another important consideration for finite element mesh during mesh coarsening is the introduction of hanging nodes are elements faces and edges, see Fig 6. It is important to maintain a 2:1 element size ratio between neighboring elements to maintain non-linear material evolution around these boundaries. Fig. 7 shows undesirable neighboring elements with excessive hanging nodes. Another side effect of excessive hanging nodes is the stiffness matrix rank degradation which results in slower convergence of iterative equation solvers generally used in finite element (FEM) simulations.Fig. 6 Coarsened elements with hanging nodes maintain 2:1 ratio.Fig.

6 Coarsened elements with hanging nodes maintain 2:1 ratio.Fig. 7 Excessive coarsening resulting in undesirable hanging node pattern.During the process of mesh refinement, a crucial step involves transferring solution variables from the old mesh to the new mesh. This transfer is essential for maintaining a continuous record of the material point's temporal evolution, and its significance becomes particularly pronounced in non-linear finite element method (FEM) simulations. In our work, we've leveraged the 2:1 ratio-keeping principle to streamline this mapping process, significantly enhancing its speed and efficiency. This principle comes into play because, as we transition to a coarser level of1734mesh, the Gaussian points where material point calculations occur are situated in close proximity. This proximity allows us to simplify the transfer of material point values to the new coarser mesh.

This proximity allows us to simplify the transfer of material point values to the new coarser mesh. Essentially, we achieve this by directly copying material point values from the nearest neighboring old material points. This seamless association of nodes, critical for maintaining the continuity of solution variables, is efficiently facilitated through the use of the octree data structure. The octree data structure allows for an easy establishment of parent-child relationships between the old and new elements, ensuring that the material point values are accurately transferred, and thus, enabling the preservation of essential temporal information throughout the mesh refinement process., see Fig. 8Fig. 8 Material point association between old and new meshExampleTo demonstrate the above discussed considerations, we use an example to showcase the various aspects of achieving a coarse mesh which keeps the required accuracy at the desired location in the geometry. In Fig.

In Fig. 8 we take a typical geometry built using additive manufacturing and get a fine mesh representation of the domain. During the layer-by-layer build process various coarsening can be started after building a certain number of layers to reach a steady state in the bottom layers. In this case we started coarsening at layer 50 and did a couple of more coarsening steps at layer 60 and layer 70. We can see in Fig. 9 that the total number of elements reduces rapidly on the first coarsening and only reduces slightly on the second and third coarsening on those given layers. So, it is important to choose the start layer for coarsening and the mesh coarsening frequency. This example also demonstrates the earlier discussed aspects such as the 2:1 neighborhood, boundary fine mesh and top layer fine mesh arrangement.1735Fig.

This example also demonstrates the earlier discussed aspects such as the 2:1 neighborhood, boundary fine mesh and top layer fine mesh arrangement.1735Fig. 8 AM geometry and its voxelized representation with 515,799 elements(a) First coarsening at layer 50 – 408,846(b) Next coarsening at layer 60 – 388,770 elements.(c) Next coarsening at layer 70 – 372,243 elements.Fig. 9 Mesh coarsening at layer 50 and subsequent coarsening at layers 60 and 70 of above example geometry.ConclusionsIn summary, the field of additive manufacturing and simulation has seen remarkable progress, with layer-by-layer inherent strain and coupled thermo-mechanical simulations gaining widespread recognition. Central to the success of these simulations is the adoption of voxel-based meshes, which excel in accurately discretizing complex geometries.

Central to the success of these simulations is the adoption of voxel-based meshes, which excel in accurately discretizing complex geometries. The seamless integration of voxel meshes into octree data structures further enhances their versatility, enabling efficient mesh manipulation to achieve precise resolution where it matters most. The localized nature of1736phenomena in many additive processes underscores the significance of tailored mesh refinement. Octree-based voxel meshes prove indispensable, offering computational efficiencies and streamlined data storage solutions. The incorporation of a parent-child relationship between old and new elements effectively addresses non-linear material evolution, ensuring the fidelity of simulations to real-world dynamics. The concept of geometry-specific refinement, achieved through strategic buffer elements, enhances the precision of simulations.

Minimal overhang length is selected as the objective function to optimize the build direction. The thickness is adjusted to meet allowable overhang length and allowable cups height. The input mesh is first decomposed into non-uniform thickness segment meshes and then each segment is cut into uniform thickness slices. The output slices consist of split slices between two adjacent segment meshes and inner slices for each segment mesh. Examples and analyses confirm the feasibility and effectiveness.IntroductionAdditive manufacturing (AM) is a set of digital manufacturing technologies integrated CAD (Computer Aided Design), CAM (Computer Aided Manufacturing) and CNC (Computer Numerical Control). Different from the past, more work is now focused on metal than plastic and various kinds of direct fabrication processes such as SMD (Shaped Metal Deposition), LENS (Laser Engineered Net Shaping) and SLM (Selective Laser Sintering) have been sprung up.

In recent years, metallic component direct fabrication without the usage of sacrificial support structure has been one of hot topics in the AM field. However, it is usually necessary to add extra support for building overhang component in the conventional 2.5-axis manufacturing platform. The support will result in the following problems: 1. Design aspect. There is no uniform standard to guide the support design and it ishard to design support for components with high geometry complexity.2. Fabrication aspect. The material of support is usually different from the material of the component. The manufacturing efficiency will be reduced because of t the material transformation. In addition, it will waste the material and improve cost because the support is sacrificial.3.

The manufacturing efficiency will be reduced because of t the material transformation. In addition, it will waste the material and improve cost because the support is sacrificial.3. The remove of the support is a time-consuming process and the quality of the split surface is poor.The emergence of multi-axis manufacturing system provides the chance for metallic component direct production without the usage of support. It always keeps the previous built piece as the support of the following building piece by changing the build direction. The component is first decomposed into several sub-volumes and then each sub-volume is created along a fixed direction. Another factor is the effect of surface tension. It makes it possible to fabricate the volume with micro-overhang without the usage of the support.Therefore, it is urgent to provide a new slicing method to meet the demands of multi-axis manufacturing system.

Different the conventional slicing methods, the cruces of multi-axis slicing involve the build direction and the segmentation. Singh [1] and Sundaram [2] used isoclines to decompose the CAD model into two pieces. One is manufacturable and another is hard to make. However, it is necessary to specify the reference build direction and is hard to use for triangular mesh because of discretization error. Similar the visibility map, Singh and Dutta [1] proposed the concept of build map with respect to the overhang angle by using Gauss map. The build map is the set of all feasible build directions. On the basis of the build map, the component is decomposed and the optimal direction is computed in accordance with different targets such as minimal stair-stepping error, minimal volume error and maximal efficiency. However, it is difficult to compute the build map because it refers to complex intersection operations involving the intersection between the half-space and the sphere.

However, it is difficult to compute the build map because it refers to complex intersection operations involving the intersection between the half-space and the sphere. Therefore, it is hard to implement. Zhang and Liou [3] defined the overhang angle by using the estimated tangent vector. The build direction is optimized to meet minimal spherical crown which encloses Gauss maps of all sampling tangent vector. The computation is incremental and complex. In addition, it will lead to unexpected results when the sampling tangent vector is few.Central axis, also known as skeleton, is the 1-dimesion abstract representation for a 1-dimesion object. It has been extensively used for image dissection and retrieval [4-5]. Similar the central axis, the concept of centroid axis [6-8] is proposed and used for overhang part decomposition. The centroid axis is the loci of the centers of all slices and guides the feasible build direction. The advantage is simple and easy to implement.

The centroid axis is the loci of the centers of all slices and guides the feasible build direction. The advantage is simple and easy to implement. However, the estimating precision is closely related to the slice thickness. It is hard to balance the efficiency and the precision. In addition, the estimating centroid axis will deviate from the actual skeleton when the surface has sharp corner. In this case, it will be failed. Literature [8] provided a improved version by comparing the areas between two adjacent slices and defined several potential slicing directions. It can only figure out the case of orthogonal sharp corner.five-axis manufacturing system integrated material additive and subtractive processes, can be used to create overhang metallic components directly without the usage of sacrificial support structure. Fig.1 illustrates the fabrication course of an overhang component.

Fig.1 illustrates the fabrication course of an overhang component. The component is decomposed into five pieces each of which has micro-overhang and can be directly built. The previous piece will turn into the support for the following piece by rotating the A-axis.

The former is applied to the input mesh to generate a series of non-uniform segment meshes, each of which can be fabricated directly without the usage of support. The latter is served for each sub-mesh to generate a series of slices, each of which with the same thickness is uniform. Although both of them are implemented by a plane intersection with a mesh, the direction of the plane for the segmentation is changeable and that for the slicing is constant. It is well known that the result of the plane intersection with the mesh is a polygonal contour. To differentiate them, the plane and the contour for the segmentation are called segment plane and segment contour and they for the slicing are called slice plane and slice contour.As shown in Fig.2, given a triangular mesh M, the input arguments are listed in Table 1 and the output is a set of segment data (see Fig.3).

Each segment data consists of a coordinate system, a segment mesh, a segment under contour and a group of slice contours. The first three items are generated from the segmentation and the last item is generated from the slicing.The segment plane is determined by the centroid of the trial segment contour and the optimum build direction. To segment the mesh in a unified way, coordinate transformation is first applied to the mesh and then a zero-height segment is carried out. The objective coordinate system takes the optimum build direction as the Z-axis and the centroid as the origin. Zero-height segment decomposes the mesh into twopieces. One is called under mesh with segment plane as the top and another is called upper mesh with segment plane as the bottom. The segmentation includes initial segmentation and overhang segmentation. The former generates the initial build direction and the target mesh for overhang segmentation.

The segmentation includes initial segmentation and overhang segmentation. The former generates the initial build direction and the target mesh for overhang segmentation. The latter optimizes the build direction and segment thickness and generates segment plane and segment data. Each segment mesh is formed by two adjacent segmentations. The previous provides the under boundary and the next provides the upper boundary. The under boundary of the first segment mesh is determined by the initial segmentation and the upper boundary of the last segment mesh is determined by the input mesh. Suppose the input mesh is decomposed into (cid:1866) segment meshes.

As shown in Fig.4, let (cid:1848) be a point on the top contour (cid:1829)(cid:3047). If (cid:1848) (cid:1489) (cid:1827)(cid:3029), it is called an overhang point and is denoted by (cid:1848) (cid:1488) (cid:1841); otherwise, it is called a non-overhang point and is written by (cid:1848) (cid:1489) (cid:1841).

It converts the minimal optimization into the maximal optimization with max (cid:2187)(cid:3021)(cid:1845)(cid:2187) as the objective function and (cid:2187)(cid:3021)(cid:2187) (cid:3404) 1 as the constraint condition. By using Lagrange multiplier approach, we have (cid:1845)(cid:2187) (cid:3404) λ(cid:2187) . The eigenvector with the maximum eigenvalue for the sparse matrix (cid:1845) is our target. The sparse matrix (cid:1845) is estimated by the covariance matrix in the usage of PCA. When the number of points on the segment contour is few, there will be a greater estimation error between the sparse matrix and the covariance matrix. In this case, the confidence of the build direction will be reduced. Therefore, we must provide enough points to better estimate the sparse matrix.

In this case, the confidence of the build direction will be reduced. Therefore, we must provide enough points to better estimate the sparse matrix. One efficient approach is the usage of high-precision input mesh.Fig.4 Overhang length between two adjacent layersFig.5 Overhang length and overhang angle for a single layerFig.6 Gauss map of build directionsSegment thickness optimizationThe segmentation consists of the initial segmentation and the overhang segmentation. For the former, the segment plane is parallel to the bottom of the AABB (Axis-aligned bounding box) of the input mesh and the segment thickness is the initial segment height. For the latter, the segment thickness is optimized to meet allowable overhang length and allowable cusp height. The concept of cusp height was proposed by Dolenc and Makela [14] and used to measure the stair-stepping error. For an overhang component, the cusp height also can be used to measure the missingvolume.

Therefore, the segment thickness under allowable cups height is formulated by:(16) (17) Except for allowable cups height, the segment thickness must be constrained by allowable overhang length. Suppose that there is a linear relationship between the segment thickness and the overhang length.

The bend is decomposed into eight segment meshes. Fig.7d illustrates the output slices. Except for the last segment mesh, each segment mesh has one segment contour and two slice contours. The last segment mesh only has one slice contour.Fig.8 illustrates the overhang measurement comparison between the build direction unchanged and changed. The overhang measurement includes overhang length and overhang angle. The overhang angle and the overhang length for each point on the segment contour are calculated by Eq.(6) and Eq.(4). The allowable overhang length is 0.625mm which can be computed by Eq.(19). From Fig.8a, the overhang angle is significantly reduced and the maximum is reduced from 12.66° to 3.85°. From Fig.8b, the overhang length is reduced from 0.71mm to 0.60mm which less than 0.625mm.

The corresponding input arguments are listed in Table 3. The bend is decomposed into six segment meshes. The segment thickness is small where the curvature is big and is big where the curvature is small. Fig.9c illustrates the adjustment of segment thickness. The pre-segment thickness, optimized thickness and actual thickness are computed by Eq.(16), (17) and (18) respectively. Build directions of the first segment mesh and the last segment mesh are parallel and their overhang lengths are less than the allowable value, therefore their segment thicknesses are the maximum allowable segment thickness. The thickness of the 2nd-4th segment meshes are less than the minimum allowable segment thickness. The thickness of the 5th segment mesh is 5.43 which is between the minimum segment thickness and the maximum segment thickness.

The thickness of the 5th segment mesh is 5.43 which is between the minimum segment thickness and the maximum segment thickness. Fig.10 illustrates a more complex application example..25 5 2Table 3 Input arguments for Fig.9/mmAllowable cusp height Maximum segment thickness Initial segment heightcmax tmax sh1.2 Allowable overhang length 9 Minimum segment thickness 0Slice thicknesssolmax tmin stAdjustment of segment thickness1110Trial segment thickness Optimum thickness Actual thickness9)m m8(s s e n k c h ti76t n e m g e S54320123456Segment index(a)segment meshes; (b)output slices; (c) adjustment of segment thickness Fig.9 Adaptive slicing for the bend with variable segment thicknessesFig.10 Adaptive slicing for the helixConclusionsTo realize overhang component direct fabrication using multi-axis HPDM, an adaptive slicing algorithm with variable direction and thickness is proposed in this paper.

It consists of multi-directional segmentation and single-directional slicing. The.15 5 17former is applied to the input mesh to generate a series of non-uniform segment meshes, each of which can be fabricated directly without the usage of support. The latter is served for each sub-mesh to generate a series of slices, each of which with the same thickness is uniform. The segment direction is optimized to meet minimal overhang length and the segment thickness is adjusted to meet allowable overhang length and allowable cups height. The output slices consist of split slices between two adjacent segment meshes and inner slices for each segment mesh. The algorithm has been implemented by using C++ and OpenGL in the platform of VS2010. Examples and analyses confirm its feasibility and effectiveness.AcknowledgementWe gratefully acknowledge the National Science Foundation of China (Grant Nos.51175203 and 51374113) for supporting this work.References1 Singh P, Dutta D. (2001).

aptive Slicing of Solid Freeform Fabrication ProcessesYong Seok Suh and Michael J. Wozny· Rensselaer Polytechnic InstituteTroy NY 12180AbstractThe Solid Freeform Fabrication (SFF) process significantly reduces part specific setup manufacturing lead time. This process has been primarily used in fabricating prototypes for design the major impact of this process on the future of visualization and verification. However, manufacturing technology would be the possibility of fabricating functional parts for end use. One of the obstacles to this goal is the insufficient accuracy of the final physical part produced by the process. From the software point of view, the major sources of the inaccuracy come from the inappropriate data transfer format and the 3D aliasing' or Stair-stepping' problem.The '3D aliasing' problem can be reduced by adapting the layer thickness to the geometry of the part.

In this paper, the procedure of adaptive slicing from the exact representation of the part model is described. This will improve part accuracy and minimize building time especially for the parts with highly curved surfaces. The procedures are implemented and a comparison to the conventional uniform layer thickness method will be discussed.1. INTRODUCTIONThe primary application of the SFF process has been to fabricate prototypes of new designs for their quick visualization and verification. For this purpose only, high part accuracy is not generally required. However, as the process is improved, fabrication of final functional parts is becoming more interesting. To fabricate functional parts, two major obstacles to the goal must be overcome. The first is the fact that most of the processes build parts made of limited special materials only.

To fabricate functional parts, two major obstacles to the goal must be overcome. The first is the fact that most of the processes build parts made of limited special materials only. in general, The second is an accuracy problem: the process has a relatively loose tolerance, compared to others such as NC machining[1].important consideration. The major sources of inaccuracy in the process are part shrinkage, approximation of CAD models during data transfer with CAD systems, and the stair-stepping or 3D aliasing problem[2]. A tessellated model is transferred to the SFF process from a CAD system via a file format that stores an unordered list of vertices which comprise triangular facets. Even though a part is modeled with exact geometries in a CAD system, valuable geometric and topological information is lost during the data transfer[3]. The finite thickness of layers of physical material produces 3D aliasing effects that cause the inaccuracy of the final physical part[4].

The finite thickness of layers of physical material produces 3D aliasing effects that cause the inaccuracy of the final physical part[4]. This is illustrated in Figure 1.In building a part using the SFF process, part accuracy is usually the mostCurrently at National Institute of Standards and Technology404i/+"'~ ---') ..-..,.,.,~~~--I.._ . /;:.._..?/'i woo lII Ij !..........................................................................................................................................................................mFigure 1. 3D aliasing effect It is not possible to eliminate 3D aliasing completely because of the process characteristics. However, it can be reduced by varying the thickness of the slices.

3D aliasing effect It is not possible to eliminate 3D aliasing completely because of the process characteristics. However, it can be reduced by varying the thickness of the slices. The thickness must be adaptive to the shape of the surface of the part in order to guarantee that the deviation between the final part surface and the ideal part surface lies within a certain tolerance limit.In this research, the issue of accuracy is concerned primarily with focusing on reducing the 3D aliasing problem by optimizing the layer thickness. The exact model is transferred from CAD systems through high-level data exchange standards such as STEPIPDES[5].2. OVERALL SYSTEM ARCHITECTUREFor a fully automated SFF process, a system architecture shown in Figure 2 is currently being developed and tested at Rensselaer Polytechnic Institute. A solid modeling kernel is built as a part of the process software to store model information and perform geometric reasoning.

~c:::IAuIuCAD. . Database --.Design SystemData Exchange Figure 2. System ArchitectureSFF Machine._-----~~~--------_._._----------------.~~----_._-----SFF System405..3. ADAPTIVE SLICING PROCEDUREThe purpose of the adaptive slicing is to fabricate geometrically accurate parts. The reference of the part accuracy is the solid model with mathematically defined surfaces. Many attempts have been made successfully to manufacture parts directly from CAD models: the automatic NC machine tool path generation is an example of the attempts.The 3D aliasing causes the deviation of the part surface from the true surface definition, and the deviation is defined by the maximum gap measured in the normal direction of the surface as shown in Figure 3. The cusp height tolerance is set by designers or SFF process operators based on the functionality of the part to be built. Each surface of a part may be set with different tolerance values according to its functionality.

Each surface of a part may be set with different tolerance values according to its functionality. The following subsections will discuss the procedure of adaptive slicing in sequence based on the assumption that a cusp height tolerance and a range of allowable layer thickness are given with a properly orientedpart model.3.1 Finding Slicing Sub-regionsSome special features must be considered when slicing a part for accuracy. Those features include flat areas, peak points and edges, and horizontal edges and the slicing plane must pass through those features for accurate fabrication. The inaccuracies caused by ignoring those features are illustrated in Figure 4. These features will be called peak features throughout this paper."""----~....--'=:I I Layer thicknessFigure 3. Cusp heightFigure 4.

These features will be called peak features throughout this paper."""----~....--'=:I I Layer thicknessFigure 3. Cusp heightFigure 4. Inaccuracies due to peak featuresOnce all the peak features are identified in a model, the height of the features is sorted in the order of vertical direction, then the whole part can be subdivided into several sub-regions so that no peak features are within a region except its boundaries. An adaptive slicing algorithm, which will be discussed in the next section, is applied to each subregion. The peak features are identified from the vertices, edges and faces of the model, as described in the following three subsections.3.1.1 Vertex peak featuresAll the vertices are basically considered as peak features. However, two types of vertices areexcluded from the peak feature list. • topological vertex Some vertices exist only to fulfill the B-rep data structure requirements of a solid modeling system. Several examples are shown in Figure 5.

• topological vertex Some vertices exist only to fulfill the B-rep data structure requirements of a solid modeling system. Several examples are shown in Figure 5. Usually, only one or two faces are adjacent to the topological vertex.The part orientation is important because it affects the build time, dimensional acccuracy and post-processingcomplexity.406smooth vertex If a vertex lies on the intersection of surfaces that meet smoothly at the vertex, the vertex is excluded from the peak: feature list. Figure 6 shows an example of such vertices. oI Ioj.----- ..."',-- .....~~.,,,~,,," ,I I I I\ I I II II I I \,,I,,'...., ~....... ""'--'"(ll)(h)Figure 5. Topology verticesFigure 6. Smooth vertices3.1.2 Edge peak featureshorizontal edge If the geometry of an edge is a planar curve and the edge lies horizontally, the edge is considered as a peak feature as shown in the Figure 7(a).

However, if the adjacent surfaces meet smoothly at the edge as illustrated in Figure 7(b), the edge is excluded from the peak feature list.extreme pointIf an edge is not linear, the extreme points in the vertical direction must befound to be included in the peak: feature list as shown in the Figure 8.iL---ri / i[JC~(b)Figure 7. Horizontal edgesFigure 8. Extreme points3.1.3 Face peak featureshorizontal planar face The slicing plane must pass through the horizontal planar face. • extreme point Extreme points of curved surfaces in vertical direction. Once these features are found and the part is divided into regions, the procedures in thefollowing sections are applied to each region to generate a series of slicing planes.3.2 Calculation of the Next Slice HeightWithin each slicing region found at the previous section, a set of slicing planes is generated, considering a cusp height tolerance.

Let ziL and ziH be .the h~ight of th~ lower an~ higher boundaries of the ith slicing region respectively, and ZlL < zlJ < z12 < ... < Z) < ... < ZlH be an ordered sequence of height of the slicing planes. The height of a slicing plane, Zj+ J, is determined from the contours (intersection curves) generated by the previous slicing plane at the height of zj> based on the geometry of the part.The procedure is basically similar to that ofNe tool path generation for curved surfaces[8,9]. Let Cj be the contour at the height of Zj and Cj+ ] be the next contour to be found. In calculating407the next layer thickness, d = Zj+ 1 - Zj, a set of sampling points are selected along the previous contour, Cj' and the next layer thickness is calculated at each sampling point considering the geometry of the surface at that point. The minimum value is selected as the next layer thickness among those values calculated at every sampling point.

The minimum value is selected as the next layer thickness among those values calculated at every sampling point. The method for selecting a set of sample points will be discussed in the next section.3.3 Determination of the sampling pointsAssume that a sample point, Pi, on a contour curve is obtained. The next sampling point, Pi+I, can be calculated from the Figure 9 and Eq.(l). The contour curve is approximated by an arc with the radius of p=lhc where K is the curvature of the curve at the sampling point Pi. The Eq.(1) calculates an arc length to the next sampling point so that the maximum deviation, 0, is less than or equal to the cusp hr-el....igl-hl..:...t..:...to;...l..;;,;er;...an:;;.;..;;..ce::..:,..--.Pi: current sampling point Pi+l : next sampling point o : cusp height tolerance p : radius of curvature at Pi I : arc lengthFigure 9.

Determination of sampling pointsifp= 0otherwise(1)3.4 Calculation of layer thickness at a sampling pointIn this section, a method for calculating the next layer thickness at a sample point will bediscussed.dFigure 10. Calculation of next slicing plane Let p be the radius of curvature of the part surface in the vertical direction at the sampling point, P. Then the part surface can be approximated by a sphere with the radius of p as shown in Figure 10. This is a good approximation because the layer thickness is much smaller than the radius of curvature of the surfaces of most mechanical parts. The next layer thickness, d, can be tolerance, o. It is assumed that the laser calculated from the Figure 10, with the given cusp height408beam radius is compensated at the border of the part to allow for the laser line width.

It is assumed that the laser calculated from the Figure 10, with the given cusp height408beam radius is compensated at the border of the part to allow for the laser line width. The cross sectional shape of a curved line is approximated by a rectangle2.If P is infinite, the next layer thickness, d, is calculated using Eq.(2),00(= I max)if nk = 0d={8/cos8 = 8/nkotherwisewhere nk is a vertical component of the normal vector of the surface at the point P. Note that the layer thickness, d, must be bounded by the maximum and minimum layer thicknesses, lmax andImin·If r is not infinite, it is required to check the location of the sampling point, P, relative to the center of the sphere. If the point is on the northern hemisphere, the layer thickness, d, is calculated using Eq.(3).

If the point is on the northern hemisphere, the layer thickness, d, is calculated using Eq.(3). The Eq.(4) is for the case where the point is on the southern hemisphere of the sphere.d = -psin8+~p2 cos28-28p-82 d ={PCOS8-~p2 cos pcos82 8-28p-82if p2cos2 8-28p-8otherwise2> 03.5 Layer Contour Generation at Region BoundariesCare must be taken when generating contours of a layer that pass through a region boundary. All the pathological cases of surface intersection problems may occur because of the peak features. Another concern is the removal of the contours generated by downward planar faces. As pointed out by Dolenc and Makela[ 10,11], if a slicing plane passes through a down-facing flat surface, the corresponding contour area must be discarded.These problems can be eliminated by lowering the slicing plane by an infinitesimal amount which is far less than the minimum layer thickness, but greater than the tolerance value used in surface intersection procedure.4.

IMPLEMENTATION AND EXAMPLESA prototype program has been implemented that includes the features discussed in this paper. Based on the ACIS®3 solid modeling kernel, the system is built on a Sun Sparc workstation with an X-window based user interface.To compare the adaptive slicing with the uniform slicing, a sphere with the diameter of 10" is modeled and sliced with various cusp height tolerances as shown in Figure 11. The minimum and maximum layer thickness is set to be 0.001" and 0.020" respectively. The results are summarized in the table below.Adaptive SlicingCusp Height (in.) 0.006 0.007 0.008Number of Layers909 803 726Uniform Slicing Number of Layers cusp heightThick.1667 1429 1250The shape of a cured line is better approximated by a parabolic cylinder 2.

2 3 ACIS is a registered trademark of SPATIAL TECHNOLOGY Inc., Boulder, CO409(2)(3)(4)0.009 0.010 0.011 0.012 0.013 0.015669 626 593 658 548 5331112 1000 909 834 770 715With the cusp height of 0.01", the number oflayers is 626. If the uniform thickness method is used, the number of layers is between 500 and 10,000 according to the layer thickness chosen. For the uniform thickness method, the deviation error (cusp height) will vary according to the latitude of the sphere and the maximum error will occur around the north and south poles.

For the uniform thickness method, the deviation error (cusp height) will vary according to the latitude of the sphere and the maximum error will occur around the north and south poles. For the uniform layer thickness of 0.011", for example, the number of layers will be 909 and the cusp height at the angle of 80 degrees from the equator will be 0.0109", which can be achieved with only 593 layers by using the adaptive slicing method.From this experiment, we can conclude that, in many parts with highly curved surfaces, a part can be fabricated by using the adaptive slicing method, not only more accurately but also more rapidly. Figure 12 and 13 show more examples.5. CONCLUSIONThe Solid Freeform Fabrication process is a promising new manufacturing method that can fabricate a product with lot size of 1 with minimum cost. In order to serve this goal, however, further development of the process is required.

In order to serve this goal, however, further development of the process is required. From the software point ofview, the part accuracy is the major area to be tackled first. The sources of inaccuracy is identified to be from the approximate model data caused by the data exchange format and '3D aliasing' problem. In this paper, a new system architecture was proposed to exchange exact model representation with CAD systems. And adaptive slicing procedure for reducing the 3D aliasing problem was also discussed.The procedures has been implemented and tested on several example parts. The results show that the adaptive slicing assures part accuracy and minimizes the number of layers for minimal build time...., ..-;: / / ...."Figure 11. Adaptive slicing of a sphereFigure 12. Adaptive slicing of a CAM-I part410Figure 13. Adaptive slicing of a NURB surface6. REFERENCES[1] B.Hirsch and H. Muller. "StereoLithography - fields of application and factors influencing the accuracy.

APTIVE SLICING WITH THE SANDERS PROTOTYPE INKJET MODELING SYSTEMK. Unnanon, D. Cormier, and E. Sanii Department of Industrial Engineering, North Carolina State University, Raleigh, NC 27695AbstractThis paper presents one of the first known in depth studies of the Sanders Prototype inkjet modeling process. A process capability study was performed in order to determine the relationship between process parameter levels and the resulting surface roughness. The data was used to create a predictive model of surface roughness using a backpropagation neural network. Test results indicate that the network is quite effective at generalizing to new process configurations. The predictive surface roughness model is used in a newly developed inkjet modeling adaptive slicing algorithm. On a region-by-region basis, the algorithm determines the fastest machine configuration that can be used to build a part while satisfying the surface roughness requirements.

On a region-by-region basis, the algorithm determines the fastest machine configuration that can be used to build a part while satisfying the surface roughness requirements. The adaptive slicing system has been tested, and results documenting substantial time and cost savings are presented.IntroductionInkjet modeling is a relatively new rapid prototyping process that has quickly gained acceptance in the marketplace. The process is an adaptation of inkjet printing in which ink is replaced by a molten thermoplastic material. The molten thermoplastic material is plotted and solidified one layer on top of the next until the part is complete. Although the Sanders Prototype ModelMaker II system has been relatively successful in sales for a number of years [Wohlers, 2000], virtually no published studies of the process have yet appeared in the literature. To some degree, quality is process-dependent in that part quality is influenced by material properties and process parameters.

To some degree, quality is process-dependent in that part quality is influenced by material properties and process parameters. Although most SFF processes employ a layered manufacturing technique, their material and processing techniques are significantly different from one another. To understand the influences of parameters of a specific SFF technique on the quality of parts, a study must be performed using that technique.Parts built via any layered manufacturing technique are characterized by "stair stepped" surfaces that can result in a rough finish on the part. The use of very thin layers can alleviate this problem, however, this dramatically increases prototyping time and cost. In order to achieve balance between build speed and part quality, some researchers advocate the use of adaptive slicing algorithms that examine the part geometry and attempt to use the thickest possible layers in a given region that satisfy the surface roughness requirements.

One of the challenges with adaptive slicing is that it is dependent on process parameters and is therefore process specific.Adaptive slicingSlicing with constant layer thickness is common with rapid prototyping systems. However, the use of constant layer thickness can unnecessarily increase build times due to the fact that the thinnest layers needed to achieve the finest overall surface requirement are used for259the entire part. The use of variable layer thickness with adaptive slicing can help reduce build times, hence several adaptive slicing algorithms have appeared in the literature in recent years. Sreeram et al. (1995) consider the problem of generating variable thickness slices for the layered manufacturing of prismatic objects subject to a user specified cusp height constraint. They formulate an optimization problem where the number of slices is to be minimized subject to a given upper bound on the cusp height.

They formulate an optimization problem where the number of slices is to be minimized subject to a given upper bound on the cusp height. The result is an optimal orientation for the part such that the number of required slices is minimized while maintaining the specified cusp height.Kulkarni and Dutta (1995) develop a procedure for the adaptive slicing of a parametrizable algebraic surface using cusp height criterion. The acceptable cusp height is specified by the user. The slicing technique computes slices of varying thickness based on the bounding surface geometry to satisfy user-defined cusp height. The geometry is limited to parameterizable algebraic surfaces so that the exact expression for the surface curvature in the vertical direction can be extracted.Sabourin et al. (1996) present an adaptive slicing method using stepwise uniform refinement. The CAD model is first sliced uniformly into slabs of thickness equal to the maximum available fabrication thickness.

(1996) present an adaptive slicing method using stepwise uniform refinement. The CAD model is first sliced uniformly into slabs of thickness equal to the maximum available fabrication thickness. Each slab is then re-sliced uniformly as needed to maintain the desired surface accuracy. The slicing technique examines both the top and bottom of each slice to guard against sudden changes in curvature above the base of the layer.Hope et al. (1997) present an adaptive slicing procedure for improving the geometric accuracy of layered manufacturing techniques that use layers with sloping boundary surfaces. Unlike other adaptive slicing techniques, this technique uses layers with sloping boundary surfaces that closely match the shape of the required surfaces.Implementations of adaptive slicing are often based upon highly idealized circumstances. It is typically assumed that surfaces exhibit theoretically perfect stair stepping.

It is typically assumed that surfaces exhibit theoretically perfect stair stepping. According to this assumption, the roughness of a particular surface is a function of only layer thickness and surface slope. In practice, however, surface roughness is influenced by many parameters and is process- dependent. For example, the two photographs in Figure 1 show a smoother up-facing surface (left) and a rougher down facing surface (right) from parts built using the Sanders Prototype process. Both photographs reveal that the process clearly does not exhibit perfect stair stepping. The same can be said for most SFF processes.Figure 1 Edge profiles of parts built using the Sanders Prototype process260In SFF processes, surface quality of the prototypes is significantly influenced by build parameters, i.e. machine configuration and part orientation. The use of proper build parameters not only improves the surface quality but also reduces fabrication time and cost.

machine configuration and part orientation. The use of proper build parameters not only improves the surface quality but also reduces fabrication time and cost. Usually, the decision making process on the selection of the build parameters is left to machine operators, or is sometimes not even considered. Thus, quality of the prototypes depends largely on the operators’ experience and how well they select the build parameters. To enhance the part quality, it is important to have a systematic tool that can adaptively select the appropriate build parameters. Accordingly, this research has attempted to achieve this task by adding some intelligence into the slicing process so that important considerations are taken into account when the slicing process is executed.MethodologyThis section describes the methodology employed in the development of the adaptive slicing system for the Sanders Prototype inkjet modeling process. The basic slicing approach is as follows.

The basic slicing approach is as follows. For all facets in each layer, the predictive model predicts the resulting surface roughness corresponding to a given machine configuration and surface orientation. The machine configurations are tested from the fastest to the slowest. If the surface roughness requirements for all facets in the layer are satisfied, then the machine configuration is selected for that layer. The first configuration that satisfies the surface roughness requirements for a given layer is selected. The configurations are adaptively selected layer by layer from the base to the top. Finally, adjacent layers to be built by the same configuration are grouped together into ranges.In order to implement this basic slicing approach, the development process shown in Figure 2 has been adopted.

The approach consists of two major steps - the development of the surface roughness predictive model, and the application of the predictive model within the slicing algorithm to find the optimal build parameters.Development process of predictive modelFinding optimal build parametersSTL fileExperimental DesignPre-processingNeural networks trainingPredictive ModelAdaptively determine machine configurationSlice ranges with optimal configurationsFigure 2 Development process of the adaptive slicing system261Predictive model of surface roughnessIn order to make a part with acceptable surface roughness, one needs to know how a combination of parameters would affect the resulting surface roughness so that he or she can select a proper set of parameters. Experience may help one select a good set of parameters, however, this would be rather subjective. For inexperienced operators, it might be a matter of trial and error.

Experience may help one select a good set of parameters, however, this would be rather subjective. For inexperienced operators, it might be a matter of trial and error. Trial and error is a very expensive way to gain experience. A better way to do this is to develop a predictive model based on experimental results, which can be used to determine build parameters. Hence, experiments on surface roughness have been conducted, and data was used for building the predictive model.The predictive model was developed employing neural network technology. In contrast to the classic statistical approaches, neural nets require no explicit model or limiting assumptions of normality or linearity. A neural network is a powerful tool in applications where formal analysis would be extremely difficult or impossible, such as pattern recognition and nonlinear system identification.

A neural network is a powerful tool in applications where formal analysis would be extremely difficult or impossible, such as pattern recognition and nonlinear system identification. With regards to the Sanders Protoytpe inkjet modeling system, a limited number of configurations are available on the machine, hence it is virtually impossible to design a set of experiments that independently vary each process parameter. Neural networks therefore appear to be a good approach to identifying the influence of each parameter on the surface quality. Introductory treatment of neural network principles and applications can be found in Zurada (1992).For this research, a four-layer backpropagation neural network was used (i.e. one input layer, one output layer, and two hidden layers). The first step in applying a neural network to this problem was to determine the structure of the network as well as the format of the input and output nodes.

The first step in applying a neural network to this problem was to determine the structure of the network as well as the format of the input and output nodes. For this problem, the input nodes capture information about layer thickness (T), the number of walls on build (Wb) and support materials (Ws), cell sizes of build (Cb) and support materials (Cs), and the number of close-offs (N), as well as the information about workpiece surface orientation with respect to the build direction and slab milling cutter direction. For those not familiar with the Sanders Prototype process, an explanation of these parameters is available in [Sanders, 1997]. The network has a single output node that predicts what the surface roughness will be when the conditions described in the input data exist.

The network has a single output node that predicts what the surface roughness will be when the conditions described in the input data exist. The output node produces a real number in the interval [0.000, 1.000], where 0.000 represents a surface roughness of 0 min and 1.000 represents a surface roughness of 1,000 min.In order to generate data to train and test the predictive neural network, a set of preliminary process capability studies were performed. A test piece with several cubes arranged in a spiral was prototyped using numerous different configurations that are available on the Sanders Prototype ModelMaker II machine. For each configuration, a Mitutoyo surface profilometer was used to measure the surface roughness of surfaces whose slope relative to the build plane ranged from 0 to 360 degrees.In the training process, a subset of the training data was presented to the network.

The standard backpropagation algorithm was used to adjust the internal weights of the network such that the different between the predicted surface roughness and the actual surface roughness for a given set of input conditions was minimized. Upon completion of network training, the data not262present in the training set was used to test performance of the network on configurations the network had not previously been exposed to. In this manner, it was possible to test the network’s ability to generalize.Determining the machine configurationThe predictive model for surface roughness, described in the previous section, is used to determine a proper machine configuration for a layer. For each layer, all facets intersected by the slice plane have their roughness value predicted using the predictive model.

For each layer, all facets intersected by the slice plane have their roughness value predicted using the predictive model. Starting with the configuration having the fastest plotting rate, the machine parameters together with the normal direction of all the intersecting facets, one by one, are input to the predictive model. If the surface roughness value of every facet intersected by the slice plane is below its specified allowable surface roughness, then the configuration is selected for that particular layer. Otherwise, the process is repeated using the next fastest configuration. The process continues until a satisfactory configuration is found. If no existing configuration can satisfy the surface roughness requirements, then the user might have to consider modifying the surface roughness requirements. This procedure determines the configuration that should be used for each layer. Recall that the configuration includes process parameter settings in addition to layer thickness.

This procedure determines the configuration that should be used for each layer. Recall that the configuration includes process parameter settings in addition to layer thickness. The output of this algorithm determines how a part should be sliced, and for each sliced layer what the proper configuration should be.ImplementationThe implemented adaptive slicing system consists of two main modules: a surface roughness prediction module and a configuration selection module. These modules were developed separately using Visual C++ and are independent from one another.To illustrate the effect of the build parameters on the surface quality, the predictive model was used to predict surface roughness of a vase-shaped part built using different configurations. Figure 3 shows the predicted surface roughness of the part when it is built using three different configurations.

Figure 3 shows the predicted surface roughness of the part when it is built using three different configurations. Using the OpenGL graphics library, the system provides a visual representation of predicted surface roughness so that users can manually investigate the surface quality. The example shown in Figure 3 illustrates an interesting phenomenon. For a rotational part with constant surface slope relative to the build plane within any given slice, intuition suggests that the surface roughness should be equal around the circumference of the part. The Sanders Prototype inkjet process employs a slab milling operation between layers. Experimental data indicates that workpiece orientation with regards to the milling cutter (i.e. cutter entrance face versus cutter exit face) has an effect on surface roughness. Figure 3 illustrates that the network does indeed distinguish between faces that are equal in all respects except orientation relative to the slab milling cutter.

Figure 3 illustrates that the network does indeed distinguish between faces that are equal in all respects except orientation relative to the slab milling cutter. This is a perfect example of a situation where process specific models of surface roughness are needed.263Rough (1000 µ in.)Smooth (0 µ in.)Figure 3 Predicted surface roughness built by different configurationsAn Adaptive Slicing Test CaseAn octagonal part shown in Figure 4 was used to verify system performance. The part was modeled using the SolidWorks CAD system, and was exported into the binary STL format. Unfortunately, the STL format contains no information of surface roughness requirements. Therefore, a pre-processor was used to specify each surface’s roughness information. The unused 2 bytes in each facet of the STL files were used to store the surface roughness information.

Therefore, a pre-processor was used to specify each surface’s roughness information. The unused 2 bytes in each facet of the STL files were used to store the surface roughness information. The details of non-uniform surface roughness specification are available in [Cormier, 2000].360472Build Direction330(a)(b)Figure 4 An octagonal part used for verification of the slicing systemFor this test, only the surfaces on the eight sides of the octagon had their roughness values specified. The other surfaces were ignored in the configuration selecting process. All of the eight surfaces were “individually specified” to have a roughness value of not more than 300 inches. The part was adaptively sliced, and the system recommended the three slice ranges with corresponding machine configurations shown in Figure 4 (b). ModelMaker II configuration numbers 330, 472, and 360 were recommended by the system for the bottom, middle, and top slice ranges respectively.

ModelMaker II configuration numbers 330, 472, and 360 were recommended by the system for the bottom, middle, and top slice ranges respectively. To verify the system effectiveness, the octagon was built with the specified slice ranges using the Sanders Prototype ModelMaker II rapid prototyping machine. The total build time was 29.82 hours. The surface roughness values on the eight faces werem264measured across the layer ridges using a Mitutoyo surface profilometer. Figure 5 shows the actual and predicted surface roughness values of the part.218 (212)325 (268)Actual (Predicted)159 (262)203 (287)275 (258)280 (275)129 (175)163 (201)Build DirectionCutter directionFigure 5 Surface roughness values of the eight surfaces of the octagonThe results show that 7 out of 8 faces had surface roughness values less than 300 m inches. There was only one face that exceeded the maximum allowable surface roughness of 300 inches.

There was only one face that exceeded the maximum allowable surface roughness of 300 inches. The top right face in Figure 8.9 had an actual roughness of 325 m inches, which exceeds the allowable roughness by 8.3%. Without adaptive slicing, a single configuration would have been used to produce this part. Configuration 330, which is the finest configuration among the three, would have been required. The estimated build time using this configuration is 47.56 hours, as opposed to 29.82 hours that was needed for the adaptively sliced part. Thus, the adaptive slicing procedure developed as part of this research reduced build time by nearly 40% for this part, while satisfying the required surface roughness on 7 out of the 8 surfaces. The one surface that did not meet the roughness requirement was only 8.3% away from the target. At the present time, the neural network has been trained on a relatively small set of sample data due to the expense of producing test parts.

At the present time, the neural network has been trained on a relatively small set of sample data due to the expense of producing test parts. Future efforts will concentrate on building a larger set of test parts in order to further improve the predictive accuracy of the neural network.mSummary and Future WorkIn summary, this paper has described the development of an adaptive slicing system for the Sanders Prototype inkjet modeling process. The system uses a backpropagation neural network to predict surface roughness for a given build configuration and workpiece orientation. The predictive model is used within an adaptive slicing algorithm that identifies optimal slice ranges for the part.Most adaptive slicing implementations have assumed a uniform cusp height requirement that applies to all surfaces on the part. However, few parts require the same quality on every surface.

However, few parts require the same quality on every surface. Thus, the most stringent cusp height is usually applied to the entire part if a uniform cusp height requirement is used. Alternatively, with non-uniform cusp height requirements,265unnecessarily stringent cusp heights on some surfaces can be avoided. Cormier et al. (2000) have proposed a method to specify non-uniform cusp heights. The method constructs logical surfaces by detecting abrupt changes on edges of the facets. The edges with abrupt changes are considered boundary edges. All facets falling within a given boundary loop are labeled as belonging to the same surface. Each surface is graphically highlighted on the computer screen, and the user interactively specifies cusp height requirements for each surface. The specified values are saved in the unused two bytes of each facet in STL files. With non-uniform cusp height specifications, further gains in adaptive slicing efficiency can be achieved.

The specified values are saved in the unused two bytes of each facet in STL files. With non-uniform cusp height specifications, further gains in adaptive slicing efficiency can be achieved. In fact, this method has been used as a pre-processor for specifying surface roughness on critical surfaces. To date, the capability of specifying non-uniform surface roughness has not been embedded within the adaptive slicing system described in this paper. Therefore, one area of future work will be to incorporate that functionality into the adaptive slicing software.As described in this paper, the parameter selection process considers only one orientation in which parts are sliced. It is not always the case that the orientation used for modeling the part is an optimal orientation to build the part. Therefore, the part should not be simply sliced and built in the direction it was designed. Instead, the part should be built in the direction that yields acceptable quality for the lowest cost.

Therefore, the part should not be simply sliced and built in the direction it was designed. Instead, the part should be built in the direction that yields acceptable quality for the lowest cost. In finding the best orientation to build the part, the total build time of each orientation can be used as the basis for comparison. The build time is a good basis for comparison because it reflects other fabrication costs, such as material costs and operating costs. As plotting time increases, material requirements, post-processing time, and machine utilization all increase as well. The orientation with the least build time should be the best orientation to build the part. The adaptive slicing system described in this paper is also being extended to consider optimal workpiece orientation.ReferencesCormier, D., Unnanon, K. and Sanii, E.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperADAPTIVE VOXELIZATION FOR RAPID PROJECTION GENERATION IN COMPUTED AXIAL LITHOGRAPHYKevin Coulson, Joseph Toombs, Magnus Gu, Hayden TaylorDepartment of Mechanical Engineering, University of California, BerkeleyAbstractComputed axial lithography (CAL) is a tomographic additive manufacturing technology that offers exceptionally fast printing in a wide range of materials. CAL involves pre-computing a sequence of light patterns to be projected into a photopolymer. For a uniform spatial discretization of the target geometry, computational time scales inversely with the cube of the discretization pitch, which makes it challenging to exploit the full space-bandwidth product of available spatial light modulators. This work introduces an adaptive voxelization approach to reduce computational expense.

This work introduces an adaptive voxelization approach to reduce computational expense. Using one of several proposed mesh-based complexity analyses, a CAD model is recursively subdivided into stacked sub-meshes of varying voxel resolution. These complexity methods can be tailored to emphasize complexity in particular regions. Each sub-mesh is then independently voxelized before projections are generated and optimized in parallel. On a four-core CPU, this method results in a precision CAL and other voxel-based additive manufacturing computations.speedup with applications in high-2 − 6 ×IntroductionComputed axial lithography (CAL) is a photopolymer resin-based volumetric additive manufacturing (VAM) method which pre-computes a projection sequence based on an inverse computed axial tomography algorithm [1]. During CAL printing, 2D images are projected through a rotating cylindrical volume such that the cumulative energy dose is sufficient to polymerize the target geometry.

During CAL printing, 2D images are projected through a rotating cylindrical volume such that the cumulative energy dose is sufficient to polymerize the target geometry. To resolve fine details and achieve a smooth surface finish, it is desirable to use high-resolution images during projection. Currently, the 2D images are pre- computed based on a uniform spatial discretization, or voxelization, of the target geometry – referred to as the ‘standard’ approach. However, the computational time of projection generation scales inversely with the cube of the discretization pitch. This limits the accessibility of high- resolution CAL printing despite improvements in the achievable optical resolution of the system.Voxel-based methods are becoming more prominent in pre-processing for additivemanufacturing.

While the .STL file format can more precisely describe the surface of a given geometry, the simplicity of the voxel format - the 3D analog to a pixel image - enables the use of complex computational techniques to improve AM processes. Examples include lattice1255generation [2], optimized build orientation [3], energy and error optimization [4], and tomographic reconstruction itself [1]. Additionally, adaptive resolution voxelization, in which the voxel size is non-uniform, has previously been applied in feature recognition. In one application, a point cloud representation of a 3D-scanned ear is analyzed and recursively subdivided according to proximity functions which capture the nearness of every point to a set of keypoints [5].

This subdivision has the effect of emphasizing distinct surface features.Several layer-based AM processes have begun to utilize adaptive layer heights for moreefficient printing, including digital light processing [6] and fused deposition modeling [7]. These methods analyze the contour of the target geometry and calculate a set of layer heights required to maintain an error threshold based on the intrinsic surface errors introduced by a given AM process. The goal of adaptive layer height methods is to improve print speed while maintaining reasonable accuracy.In this work, an adaptive voxelization scheme is constructed to enable efficient projection sequence generation for CAL printing. This process dramatically improves computational speed while maintaining reconstruction accuracy (Figure 1). First, several mesh-based complexity methods, which are used to identify complex features of the target geometry, are introduced and analyzed.

First, several mesh-based complexity methods, which are used to identify complex features of the target geometry, are introduced and analyzed. The results of this analysis drive the recursive subdivision of a CAD model into stacked sub-meshes which are then voxelized at non-uniform resolutions. Then, the algorithms used to generate and reconstruct the final projection sequence - including a parallel processing implementation - are described. Finally, computational speed and reconstruction accuracy are evaluated with respect to the current state-of-the-art.Figure 1. Cubic curves are fit to the projection generation time data for the standard and parallel processed adaptive projection computation methods. Effective resolution for the adaptive process is the base resolution for each sub-mesh multiplied by the number of sub-meshes. For the standard process, effective resolution is the same as the resolution.

For the standard process, effective resolution is the same as the resolution. The voxel error rate, a measure of predicted over-exposure, shows no consistent trends. However, the errors for each approach are within an order of magnitude. This demonstrates the benefit of the adaptive voxelization approach – significant computational speed-up without sacrificing accuracy.1256MethodsThe process for utilizing the proposed adaptive voxelization scheme is detailed in Figure 2. The complexity analyses and projection stitching are built around the pre-existing process for projection computation developed for CAL printing.Figure 2. A process outline for the adaptive voxelization approach.Complexity Analysis and SubdivisionThe goal of an adaptive voxelization algorithm is to make more efficient use ofcomputational resources by using greater computed projection resolution for more complex sections of the target geometry.

Therefore, methods of identifying complexity in a target geometry must be developed.direction (the axis of rotation for the cylindrical resin volume), unlike some VAM technologies 𝑧𝑧 [8]. In this case, the computation associated with each horizontal row of the voxelized target is self-contained. Therefore, splitting the model using projection sequence generation while losing no data. 𝑧𝑧In its simplest physical representation, CAL has no interactions between voxels in thenormal planes will enable independent1257Complexity analysis will ultimately drive the subdivision progression in the adaptivevoxelization scheme. Figure 3 compares the layer-based complexity for several geometries using each proposed method, with the overall complexity as the integral under these curves. The simplest of the proposed complexity methods, ‘n-vertices’ and ‘vertex density’, are based on mesh properties.

The simplest of the proposed complexity methods, ‘n-vertices’ and ‘vertex density’, are based on mesh properties. These methods do not require layer-based analysis and the total complexity is calculated by counting the number of vertices and calculating the vertex density of a mesh, respectively. These methods can provide a rough estimation of the complex regions of a part. Other metrics such as a bulk-volume1-to-volume ratio and surface-area-to-volume-ratio have been demonstrated to provide similar results to the vertices-to-volume ratio [9]. However, mesh- based definitions of complexity are not without limitations. Vertex and triangle metrics are dependent on mesh export resolution and mesh refinements. Surface area and volume metrics require watertight meshes, which can be difficult to maintain after subdivision because the new faces required to cap a sub-mesh at the split plane are not often trivial to create.Figure 3. Complexity comparisons.

Complexity comparisons. Models from top to bottom: Rodin’s Thinker, the Stanford Bunny, a triple vasculature for microfluidics, and an octet truss unit cell. Methods from left to right, ‘n-vertices’, ‘vertex density’, ‘perimeter gradient’, and ‘perimeter-to-area ratio.’1 Bulk-volume is the sum of the volume of an object and the volume enclose by an object: e.g., the bulk volume of a hollow sphere is simply the volume of a solid sphere of the same radius.1258The ‘perimeter gradient’ and ‘perimeter-to-area ratio’ are layer contour-based complexity methods. Using these metrics, the complexity is calculated by taking a sum of the complexity of all layers. Despite CAL’s departure from layer-based 3D printing, slicing the model can provide insight into the relative complexity along the method uses the layer-to-layer difference in section perimeter to define regions of the geometry 𝑧𝑧 with the most significant cross-section changes as complex.

The ‘perimeter-to-area ratio’ is calculated for each layer by dividing the perimeter of each layer contour by its area. In cases where there are multiple entities represented in a layer, the summed perimeter and summed area are used to compute the ratio. The perimeter-to-area ratio is simply calculated and the basis of more rigorous measures of shape complexity, like the fractal dimension index [10]. This method ultimately assigns high complexity to regions of a geometry with relatively small cross sections and, in theory, to layers with high fractal dimensions.axis of the target. The ‘perimeter gradient’Figure 4. The subdivision process for the Thinker using the ‘n-vertices’ (top) and ‘perimeter gradient’ (bottom) methods. A binary tree labelling structure is used to track subdivisions: e.g., represents the , the subdivision process continues. While upper third quartile of the model.

A binary tree labelling structure is used to track subdivisions: e.g., represents the , the subdivision process continues. While upper third quartile of the model. Ifrepresents the top half of the model and, the complexity of a given sub-mesh, exceeds 𝑖𝑖 = 0𝑖𝑖 = 01each method subdivides the model into six sections, they emphasize complexity in different regions of the geometry.𝜇𝜇𝑖𝑖𝜀𝜀The recursive subdivision algorithm (Figure 4) used is consistent across all proposed , is defined. Then, one of , of the full geometry and acomplexity analyses. First, the maximum subdivision depth, several complexity methods is used to define the complexity, complexity threshold, uniformly across subdivisions. For example, with resultant sub-meshes is 8. This threshold ensures that each sub-mesh contains less than 1/4th of the total complexity.𝑑𝑑𝑚𝑚𝑚𝑚𝑚𝑚 μ, is set using Eq. 1.

For example, with resultant sub-meshes is 8. This threshold ensures that each sub-mesh contains less than 1/4th of the total complexity.𝑑𝑑𝑚𝑚𝑚𝑚𝑚𝑚 μ, is set using Eq. 1. This threshold is chosen to distribute the complexityset to 3, the maximum number of𝜀𝜀𝑑𝑑𝑚𝑚𝑚𝑚𝑚𝑚𝜀𝜀 =𝜇𝜇 𝑑𝑑𝑚𝑚𝑚𝑚𝑚𝑚−12 1259(1)The recursive subdivision process begins, using anormal plane at the centroid of the . Subdivision terminates whennormal plane at the centroid of the . Subdivision terminates when𝑧𝑧μ𝑖𝑖𝜀𝜀In Figure 5, the reconstruction error is compared for the standard and adaptive projection generation methods. This error is based on a comparison between the dose reconstruction and the binary target. In this binary scheme, voxels contained inside and outside of the target region are noted as ‘gel’ and ‘void’ voxels, respectively.

In this binary scheme, voxels contained inside and outside of the target region are noted as ‘gel’ and ‘void’ voxels, respectively. A threshold is set using the minimum dose received in a gel voxel of the target, and the error is defined as the ratio of void voxels which receive a greater dose than the threshold to the total voxel count. This error quantifies the predicted overcuring of the print. In Figure 5, none of the complexity analyses stands out in consistently predicting the reconstruction error, although the ‘perimeter-to-area ratio’ and ‘perimeter gradient’ methods show promising results. Even between the standard and adaptive approaches, the best-suited complexity method for prediction is inconsistent.Figure 5. In the far-right column, the linear layer-based error for each model is compared for a low-resolution standard, high- resolution standard, and mixed resolution adaptive approaches. The magnitudes of errors for each are comparable.

The magnitudes of errors for each are comparable. For the Thinker (top), the ‘perimeter gradient’ captures the adaptive error best while ‘n-vertices’ captures the standard error best. For the triple vasculature (bottom), the ‘perimeter-area ratio’ and the ‘perimeter gradient’ are the best predictions for the adaptive and standard errors, respectively. This inconsistency motivates the exploration of other complexity definitions.In some cases, the results of the complexity analyses are sub-optimal: e.g., not all thecomplex features of a target geometry are recognized by an individual analysis method. Therefore, a user-manipulated method of defining complexity is implemented [7]. This custom1260definition parametrizes the perimeter-to-area ratio, contours or entities,slice, as in Eq. 2 𝑃𝑃/𝐴𝐴slice, as in Eq.

This custom1260definition parametrizes the perimeter-to-area ratio, contours or entities,slice, as in Eq. 2 𝑃𝑃/𝐴𝐴slice, as in Eq. 2 𝑃𝑃/𝐴𝐴, and the number of distinct closed𝑁𝑁𝑒𝑒𝑒𝑒𝑒𝑒𝑖𝑖𝑒𝑒𝑖𝑖𝑒𝑒𝑒𝑒𝑧𝑧(2)αβμ𝑙𝑙𝑚𝑚𝑙𝑙𝑒𝑒𝑙𝑙 = (𝑃𝑃/𝐴𝐴)⋅ (𝑁𝑁𝑒𝑒𝑒𝑒𝑒𝑒𝑖𝑖𝑒𝑒𝑖𝑖𝑒𝑒𝑒𝑒)with parameters to match the resulting complexity curve to their own intuitive understanding of complexity. The results of this method are demonstrated in Figure 6. This method enables greater control over the complexity definition, and thus, the subdivision progression. Ultimately, these changes affect the local resolution in projection computation.andas the user-modified parameters. For any geometry, the user can adjust the𝛼𝛼𝛽𝛽Figure 6. For the Thinker, (A) and (B) represent two possible complexity definitions by changing theandparameters. For theStanford Bunny, (C) and (D) serve the same purpose.

For the Thinker, (A) and (B) represent two possible complexity definitions by changing theandparameters. For theStanford Bunny, (C) and (D) serve the same purpose. For the Triple Vasculature, complexity definition (E) results in the subdivision scheme (G), while complexity definition (F) results in subdivision scheme (H). A flexible, modular complexity definition is useful for two reasons. First, it broadens the scope of applications beyond CAL-specific complexity. Second, the CAL process is still being studied to find which features are difficult to print – this method doesn’t tie down complexity to certain feature types.𝛼𝛼𝛽𝛽Projection Computation and StitchingAfter recursive subdivision, each sub-mesh leaf has a 3D target geometry computed on aCartesian voxel basis. The discretization pitch is dependent on a base resolution, scales with the inverse of greater complexity (i.e., sections at a greater subdivision depth)., and.

The discretization pitch is dependent on a base resolution, scales with the inverse of greater complexity (i.e., sections at a greater subdivision depth)., and. This approach implements increased resolution for sections of𝑟𝑟𝑏𝑏𝑚𝑚𝑒𝑒𝑒𝑒𝑑𝑑𝑖𝑖2We can now move onto projection initialization and optimization. However, the targetsfor each leaf are centered around their own centroid, rather than the global centroid. The resulting tomographic reconstructions would thus be inconsistent between leaves. To account for this offset in projection initialization, the ASTRA toolbox [11] enables the use of a rotational axis offset. This offset is calculated inandaccording to Eq. 3, Eq. 4.𝑥𝑥 𝑑𝑑𝑖𝑖Δ𝑥𝑥 = (2 Δy = (2𝑑𝑑𝑖𝑖𝑦𝑦⋅ 𝑟𝑟𝑏𝑏𝑚𝑚𝑒𝑒𝑒𝑒)/∆𝑧𝑧 ⋅ δ𝑥𝑥𝑖𝑖 ⋅ 𝑟𝑟𝑏𝑏𝑚𝑚𝑒𝑒𝑒𝑒)/∆𝑧𝑧 ⋅ δ𝑦𝑦𝑖𝑖(3) (4)1261where global height. Further explanation of these equations is found in Figure 7 below.andare the leaf offsets from the global axis of rotation and width andis theδ𝑥𝑥δ𝑦𝑦Δ𝑧𝑧Figure 7.

Further explanation of these equations is found in Figure 7 below.andare the leaf offsets from the global axis of rotation and width andis theδ𝑥𝑥δ𝑦𝑦Δ𝑧𝑧Figure 7. Rotation axis (green and blue circles) offsets from the true centroid (red) are calculated according to Eq. 3 and Eq. 4. The ASTRA toolbox accepts pixel values for the offsets, so we first transform from the model coordinate system to the voxel . Then this coordinate system. This is achieved by calculating thefraction that the sub-mesh accounts for:in fraction is multiplied by the offset in the model coordinate system, leaving us with the axes offsets in pixels. Using ⋅ 𝑟𝑟𝑏𝑏𝑚𝑚𝑒𝑒𝑒𝑒)/𝛥𝛥𝑧𝑧 𝑧𝑧 while the top section (green) will this example, the bottom section (blue) will have an x-offset of.

Using ⋅ 𝑟𝑟𝑏𝑏𝑚𝑚𝑒𝑒𝑒𝑒)/𝛥𝛥𝑧𝑧 𝑧𝑧 while the top section (green) will this example, the bottom section (blue) will have an x-offset of . This accounts for the doubled resolution of the green section.𝑑𝑑𝑖𝑖(2𝑟𝑟𝑏𝑏𝑚𝑚𝑒𝑒𝑒𝑒 = 20have an x-offset of2(2 The projection initialization and gradient descent optimization approaches used for CAL printing [1] are then used to generate individual projection sequences. Because each leaf can be handled independently, this method is well suited for parallel processing. Parallel processing has been implemented in tomographic reconstruction in previous work [12], but to our knowledge, this approach of segmenting a target is novel.⋅ 20)/100 ⋅ 𝛿𝛿𝑥𝑥 = 0.8 ⋅ 𝛿𝛿𝑥𝑥3(2⋅ 20)/100 ⋅ 𝛿𝛿𝑥𝑥 = 1.6 ⋅ 𝛿𝛿𝑥𝑥There are several steps in order to make the individual projection sets and dosereconstructions compatible. First their relative intensities need to be scaled. Second, they need to be resized to account for the different voxel resolutions.

First their relative intensities need to be scaled. Second, they need to be resized to account for the different voxel resolutions. And last, they need to be placed into a master array to account for the , which attempts to divide ‘gel’ and ‘void’ doses is calculated. To ensure uniform 3D dose dynamic 𝑦𝑦 range between leaves of the target and avoid dose discontinuities along the projection of each leaf is scaled by scaling method has been experimentally validated by printing on the CAL system.andoffsets. In optimization, a dose threshold,𝛾𝛾𝑖𝑖𝑥𝑥axis, the optimizedwithbeing an arbitrary threshold. This intensity𝑧𝑧γ𝑖𝑖/γ0γ01262The projection stitching procedure (Algorithm 1) begins by initializing a master projection set to match the highest depth voxel resolution. Projection sets are shaped aswhereis the square voxel grid length,is the count of projection angles, and are scaled by andis the voxel grid height.

Projection sets are shaped aswhereis the square voxel grid length,is the count of projection angles, and are scaled by andis the voxel grid height. Therefore, to match the scaling of each set,�𝑛𝑛𝑚𝑚,𝑙𝑙, 𝑛𝑛θ, 𝑛𝑛𝑧𝑧� the scaling factor 𝑛𝑛𝑧𝑧𝑛𝑛𝑚𝑚,𝑙𝑙 (Eq. 5).𝑛𝑛θ𝑖𝑖 𝑛𝑛𝑚𝑚,𝑙𝑙𝑖𝑖 𝑛𝑛𝑧𝑧(5)𝑓𝑓𝑑𝑑𝑚𝑚𝑚𝑚𝑚𝑚−𝑑𝑑𝑖𝑖𝑓𝑓 = 2Figure 8. An adaptive resolution projection image for the Thinker with . From top to bottom: (A) low-resolution of 80, (B) high-resolution of 160, (C) low-resolution of 80. Despite the increased effective resolution from 80 to 100, this projection set is generated more quickly than computing the projection set at the low resolution using the standard approach.𝑟𝑟𝑏𝑏𝑚𝑚𝑒𝑒𝑒𝑒 = 20Each scaled projection set is then placed relative to the center of the master projection set.

This algorithm is modified slightly to perform dose reconstruction stitching, accounting for the(Eq.𝑥𝑥12633) and a seamless projection set calculated and resolved using non-uniform resolution (Figure 8). 𝑓𝑓3) and a seamless projection set calculated and resolved using non-uniform resolution (Figure 8). 𝑦𝑦(Eq. 4) offsets – both scaled byResultsThe computational time savings using the adaptive voxelization approach aresummarized in Figure 9. The projection generation time ratio, accounting for the cubic time complexity of calculation. This lets us use collected data at all resolutions to calculate efficiency. A low represented data were collected using a 4-core CPU., is calculated according to Eq. 6,indicates more efficient computation. All theΠ(6)Tgen 3 (𝑅𝑅effective) The adaptive parallel approach ultimately results in a processing, computation time can be reduced by up to half.

6,indicates more efficient computation. All theΠ(6)Tgen 3 (𝑅𝑅effective) The adaptive parallel approach ultimately results in a processing, computation time can be reduced by up to half. The result of this approach is adaptive resolution projection computation that can be faster than using the standard approach to compute projections for the coarsest represented resolution of the adaptively voxelized target.Π =speedup. Even without parallel2 − 6 ×Figure 9. Models from left to right: Stanford Bunny, Gyroid, Octet Truss, Slot Cylinder, Sphere, Test Box, Thinker, and Triple Vasculature. With no exceptions, the adaptive method outperforms the standard approach. The addition of parallel processing to the adaptive approach further improves performance.DiscussionWhen comparing the proposed complexity analysis methods, the layer contour-based metrics align well with an intuitive understanding of complexity.

While mesh property-based1264metrics are simple, they provide less insight into the how the geometry of a target evolves along its -axis. Regardless, neither the ‘perimeter gradient’ nor ‘perimeter-to-area ratio’ methods can capture all forms of complexity. However, the user-manipulated complexity demonstrates that parametrized combinations of layer contour-based metrics can be tuned to detect complex regions of a target geometry.𝑧𝑧The complexity analysis dictates the subdivision scheme of the mesh, generating a binary tree of sub-mesh leaves. The proposed method of normalizing, scaling, and inserting independent projections and dose reconstructions from each leaf into master arrays enables the use of parallel processing while maintaining the compatibility of the projection sets. This approach is built around the existing approach for projection computation for CAL and provides a simple and robust solution.6.84.8232.16Figure 10.

This approach is built around the existing approach for projection computation for CAL and provides a simple and robust solution.6.84.8232.16Figure 10. Optimization time improvements using the non-parallel adaptive process. The adaptively voxelized and generated 𝑇𝑇𝑔𝑔𝑒𝑒𝑒𝑒,seconds, and thus an effective resolution of 100.projection set has five leaves with𝑟𝑟𝑏𝑏𝑚𝑚𝑒𝑒𝑒𝑒 = 20When compared to the standard approach, the optimization speed-up is significant for theadaptive approach. As noted above, in some cases adaptive resolution projection computation can be faster than using the standard approach to compute projection for even the coarsest represented resolution of the adaptively voxelized target (Figure 10). This result is surprising and counterintuitive, and can be explained by differences in the optimization procedure. After projections are generated, they are optimized using a gradient descent algorithm which utilizes a user-defined error threshold exit condition.

After projections are generated, they are optimized using a gradient descent algorithm which utilizes a user-defined error threshold exit condition. Because the range of gradient descent optimization algorithm operates is reduced when looking at smaller sections of the part, the solution converges – and thus falls below the error threshold – much more quickly for independent leaves than for the target as a whole. Another possible explanation is the tighter bounding boxes that can be created for independent leaves. Rather than back projecting each slice using the – the maximum cross section size – of the entire geometry, sections with smallerpositions over which the𝑧𝑧𝑖𝑖 𝑛𝑛𝑚𝑚,𝑙𝑙are calculated over a more economic domain. Additionally, parallel computing can𝑛𝑛𝑚𝑚,𝑙𝑙1265be used to optimize the solution for each leaf independently.

Additionally, parallel computing can𝑛𝑛𝑚𝑚,𝑙𝑙1265be used to optimize the solution for each leaf independently. This particular implementation of parallel processing for tomographic reconstruction is novel and provides a simple method to subdivide calculations using a parallel beam projector.ConclusionThe adaptive voxelization approach demonstrated in this work can support rapid generation of high-fidelity projection sequences for the CAL 3D printing system without sacrificing accuracy. Computational resources can be allocated towards regions of the target geometry which are identified as complex by using a non-uniform discretization pitch. The complexity analyses, subdivision method, and adaptive approach developed could be applied to other AM processes that use a Cartesian voxel basis in pre-computing.

The complexity analyses, subdivision method, and adaptive approach developed could be applied to other AM processes that use a Cartesian voxel basis in pre-computing. Future work could include an in-depth investigation into variations of the proposed complexity analysis methods, expanding on the hybrid modular approach presented in this work and potentially including the use of the fractal dimension index as a metric. Additionally, although this method enables the use of higher voxel resolutions, the voxel representation referenced as the ‘target geometry’ throughout this work does not truly represent the mesh. More future research into a mesh-based, ray trace-driven reconstruction algorithm could further improve the precision of projections used for CAL printing.

More future research into a mesh-based, ray trace-driven reconstruction algorithm could further improve the precision of projections used for CAL printing. Lastly, the development of a 3D octree subdivision structure, rather than the 1D z-slice method discussed in this work, could lead to more detailed feature recognition and subdivision, and even greater efficiency in computational resource allocation.1266References[1] B. E. Kelly, I. Bhattacharya, H. Heidari, M. Shusteff, C. M. Spadaccini and H. K. Taylor, "Volumetricadditive manufacturing via tomographic reconstruction," Sciene, pp. 1075-1079, 2019.[2] A. O. Aremu, J. J. Brennan-Craddock, A. Panesar, I. A. Ashcroft, R. M. Hague, R. D. Wildman and C. Truck, "A voxel-based method of constructing and skinning conformal and functionally graded lattice structures suitable for additive manufacturing," Additive Manufacturing, vol. 13, pp. 1-13, 2017.[3] P. Das, K. Mhapsekar, S. Chowdhury, R. Sman and S.

13, pp. 1-13, 2017.[3] P. Das, K. Mhapsekar, S. Chowdhury, R. Sman and S. Anand, "Selection of build orientation foroptimal support structures and minimum part errors in additive manufacturing," Computer-Aided Design and Applications, pp. 1-13, 2017.[4] R. Paul and S. Anand, "A combined energy and error optimization method for metal powder basedadditive manufacturing processes," Rapid Prototyping Journal, vol. 21, no. 3, pp. 301-312, 2015.[5] S. Cadavid, S. Fathy, J. Zhou and M. Abdel-Mottaleb, "An adaptive resolution voxelizationframework for 3D ear recognition," 2011 International Joint Conference on Biometrics (IJCB), pp. 1- 6, 2011.[6] J. S. Nielsen, D. B. Pederson, Y. Zhang and H. N. Hansen, "Adaptive layer height during DLPmaterials processing," Proceedings of the 2nd International Conference on Progress in Additive Manufacturing, pp. 246-251, 2016.[7] F. Wasserfakk, N. Hendrich and J.

of Mechanical Engineering, University of California, Berkeley, CA 94720, USA 4Department of Mechanical Engineering, University of Rochester, Rochester, NY 14627, USA 5Dept. of Mechanical Engineering, Massachusetts Institute of Technology, Cambridge, MA 02139, USACorresponding author: Email: shusteff1@llnl.gov Tel.: +1-925-423-0733AbstractAs additive manufacturing (AM) technologies advance and mature, the geometric constraints imposed by fabricating 2D planar layers become increasingly important to overcome. In the realm of light-driven AM fabrication, holography provides a promising avenue toward true 3D structures. Being capable of recording and reconstructing 3D information, holographic shaping of the light field can enable direct 3D fabrication in photopolymer resins.

Being capable of recording and reconstructing 3D information, holographic shaping of the light field can enable direct 3D fabrication in photopolymer resins. We have conceptualized, designed, and built a prototype holographic additive micromanufacturing system, incorporating a liquid-crystal-on-silicon (LCoS) spatial light modulator (SLM) to redirect light energy at the build volume by spatial control of the phase distribution. Here we report the system design, design parameter trade-offs relevant for producing 3D structures, and initial fabrication results.IntroductionThe conceptual framework of additive manufacturing (AM) is to make things in a bottom-up fashion, adding material to create a part or structure, rather than carving, machining, or cutting it away from a larger piece.

[1] This approach has held out the promise of efficient, rapid, and flexible fabrication of many types of three-dimensional (3D) structures and architected materials, with the important benefits of reducing wasted material, expanding geometric flexibility, and allowing the fabrication of engineered materials. Indeed, many of these advantages are beginning to be explored and demonstrated, including engineered biomimetic structures [2, 3], and engineered metamaterials with extraordinary properties [4]–[7].Every AM process can be considered to employ some “unit operation” which is repeated to build up a complex part. Considering the range of such unit operations, the most serial, lowest-throughput processes deposit or photopolymerize one volume element (or “voxel”) at a time, which can be referred to as “zero-dimensional-plus” or 0D+ fabrication (a voxel is a zero- dimensional element).

This includes technologies such as conventional stereolithography (SL) [8, 9], metal powder-bed-based laser fusing systems [10] as well as the familiar fused deposition1183modeling (FDM) method used by desktop 3D printers writing with extruded plastic. Although advanced 0D+ technologies like direct laser writing (DLW) [11] and direct ink writing (DIW) [12] offer the capability of writing arbitrary paths in 3D space, rather than being constrained to planar layers, these are still fundamentally serial processes, and inherently low-throughput. Technologies based on layer-at-once rather than voxel-at-once unit operations yield approaches that can be called “2D+”. One example is projection micro-stereolithography (PµSL) [13], in which an LED-illuminated microdisplay is used to project full 2D images to cross-link layers of photopolymer resin. This enables significant gains in build speed.

This enables significant gains in build speed. Nevertheless, this and other 2D+ approaches still carry with them some limitations inherent to building 2D layers. For instance, some geometries such as overhangs are difficult or impossible to reproduce accurately.AM processes with “volume-at-once” unit operations, to produce complex 3D structures in a single step have remained a major challenge. Some promising recent reports include the use of multi-beam interference to generate periodic 3D lattice structures [14]–[17], and to modulate such lattices with aperiodic features [18, 19]. However, volume-at-once fabrication of aperiodic geometries that overcome the geometric limitations of these lattices has not previously been demonstrated. Additionally, in 2D+ layer-by-layer approaches, resin surface settling and planarization time is one of the slowest time-limiting steps, so overcoming this bottleneck is highly desirable.

Additionally, in 2D+ layer-by-layer approaches, resin surface settling and planarization time is one of the slowest time-limiting steps, so overcoming this bottleneck is highly desirable. Ideally, this would be accomplished by fabricating away from the resin surface, within the bulk of the fluid. A secondary benefit of working within the bulk is avoiding oxygen concentration gradients associated with the fluid-air interface. Diffused oxygen plays an important role in photochemical kinetics, and proximity to a surface incurs additional complexity for properly controlling curing rates and geometry.We are therefore interested in achieving an AM process capable of depositing or solidifying a complex volume as its unit operation, providing an additional leap forward in fabrication speed, and further reducing the limitations on the geometry of fabricated parts.

In this work, we extend the polymer resin-based AM paradigm into the volume-at-once realm, providing the first demonstration of generating complex 3D structures in photopolymer as a unit operation.The enabling technology is digital holography, a powerful paradigm for both recording and reconstructing 3D geometrical information. Using a diffractive element to phase-modulate a coherent laser beam, multiple sub-images of a holographically-shaped light field are directed such that they intersect within the resin volume. The light distribution is designed to deliver energy maxima to the desired regions of the resin that are to be cured. The part forms in situ in ~10 seconds, without requiring any support structure, nor layer-by-layer curing. Our initial results produce millimeter-scale parts with ~100 µm resolution.

The part forms in situ in ~10 seconds, without requiring any support structure, nor layer-by-layer curing. Our initial results produce millimeter-scale parts with ~100 µm resolution. The results demonstrated here are suggestive of additional possibilities for this approach, and we discuss them at the conclusion of this manuscript.System Configuration and OperationA generalized light-directed AM system comprises four key subsystems, which are (1) the light source and beam conditioning, (2) pattern formation, (3) pattern projection/delivery, and (4) the photopolymer resin. While the parameter space for a holographic lithography system is1184Figure 1: Optical configuration of the holographic 3D fabrication system. Lens pairs L1-L2 and L3-L4 are telescope pairs that size the beam correctly to impinge on the SLM, with the pinhole (PH) providing spatial filtering.

Lens pairs L1-L2 and L3-L4 are telescope pairs that size the beam correctly to impinge on the SLM, with the pinhole (PH) providing spatial filtering. The Fourier-transforming lens (FTL, f=250 mm) projects the phase hologram displayed on the SLM, forming the desired intensity distribution at the hologram plane (HP). This plane is displaced from the FTL natural focal plane (where the beam block BB is placed) by phase curvature added to the phase mask of the SLM, whereas undiffracted light is removed by the BB pattern. Lenses L5 and L7 are an image relay telescope pair, projecting the 3D intensity pattern into the resin with the help of exposure mirror MEX. In a 3-beam configuration, a second mirror under the cuvette (out of plane) directs a third sub-image into the cuvette from a third orthogonal direction.

In a 3-beam configuration, a second mirror under the cuvette (out of plane) directs a third sub-image into the cuvette from a third orthogonal direction. The microscope objective MO and mirror M3 aid in monitoring image quality and exposure testing at reduced magnification when M3 is replaced by a sample.very broad [20, 21], two key considerations drive the overall technology choices and system architecture in this work. The first is that holographic shaping of the light field requires a source with a high degree of spatial and temporal coherence, for useful diffraction and interference. This system is therefore designed around a narrow spectral linewidth single-mode (TEM00) laser source. The second design consideration is that holographic beam shaping is most robust and straightforward with a phase-only spatial light modulator (SLM).

The second design consideration is that holographic beam shaping is most robust and straightforward with a phase-only spatial light modulator (SLM). With an appropriate phase pattern applied to this dynamic element the light field is Fourier-transformed by the projection optics to produce the desired intensity distribution at the build volume, photocuring the resin into the desired pattern.In the current implementation, the 3D intensity pattern is produced by folding sub- regions of the projected light field using 45-degree mirrors placed in proximity to the final build volume. These sub-regions enter the build volume at right angles to one another, effectively forming intersecting beams, each of which carries an arbitrary image pattern. The intersection of1185these patterns results in a complex 3D pattern with controllable areas of high intensity, and with appropriately chosen exposure parameters the 3D structure is formed.

The intersection of1185these patterns results in a complex 3D pattern with controllable areas of high intensity, and with appropriately chosen exposure parameters the 3D structure is formed. This approach is one of several possible paths to obtaining complex 3D light patterns, and we address alternative configurations in the Discussion section. The overall optical layout is shown in Fig. 1, and key subsystems are described in greater detail below.Laser Source and Beam ConditioningThe holographic principle allows the recording and reconstruction of 3D geometric information by capturing the amplitude and phase information contained within a light field. Holographic reconstruction requires the use of light waves capable of interference, which implies a single-frequency laser for temporal coherence. In addition, the light delivered to the SLM must ideally consist of flat phase-front plane waves, which implies spatial coherence.

In addition, the light delivered to the SLM must ideally consist of flat phase-front plane waves, which implies spatial coherence. In this work, we use a Coherent Verdi V-6 532 nm DPSS laser, specified by the manufacturer to have a linewidth of < 5 MHz (<0.000005 nm, corresponding to a coherence length of >50 m). The output beam is slightly expanded and spatially filtered through a 25 µm pinhole, to improve the circularity and spatial coherence, then expanded again such that it reaches the SLM as a Gaussian beam with a 1/e2 diameter of approximately 20 mm. This is sized to approximately match the SLM diagonal, providing a balance in the trade-off between illumination uniformity, and efficient use of laser energy.It is worth noting that the laser’s minimal coherence length requirement is much shorter than that of the Verdi DPSS laser in our system: it must only exceed the path length differences between any light paths in the system that must interfere.

In our configuration, this is only ~1 cm. However, additional coherence improves the contrast of reconstructed intensity fields at the build volume, thus benefitting lithographic fidelity; this remains to be rigorously characterized.Pattern Formation: Computer-Generated HologramsThe digital phase mask required to produce a desired intensity distribution at the build volume is referred to as a computer-generated hologram (CGH). To imprint the phase mask onto the incident laser light we employ a liquid-crystal-on-silicon (LCoS) pixelated phase-only SLMFigure 2: Holographic projection of a resolution test pattern, demonstrating the improved noise performance from averaging multiple CGHs generated to produce the same intensity pattern, initialized with a different random phase for each CGH. The simulated projection at left shows a Fourier transform of a single CGH in MATLAB, and the middle panel is the same CGH projected through our optical system and recorded with a CCD camera.

The simulated projection at left shows a Fourier transform of a single CGH in MATLAB, and the middle panel is the same CGH projected through our optical system and recorded with a CCD camera. The panel at right averages 10 separately recorded CGHs.1186(Holoeye GmbH, PLUTO VIS). The use of a phase-only SLM for holographic light projection is well established, with various investigators working to calibrate [22], linearize [23], and otherwise characterize the LC behavior [24] for reliable operation. In the 1920×1080 pixel microdisplay used for this work, each 8µm square pixel imposes a phase delay between 0 and approximately 2.1π, controlled with 8-bit image data (256 gray levels), using a linear phase lookup table provided by the manufacturer.The CGH is iteratively calculated by the well-known Gerchberg-Saxton (G-S) algorithm [25] to result in the desired light intensity distribution at the build volume.

Many other methods for calculating phase holograms have been described [26]–[29], and some of these result in substantially higher-quality images, but these often impose much greater computational costs, or some may only be used with limited geometries. The simplicity and speed of the G-S approach, and in particular its generality when producing unconstrained and arbitrary image patterns, make it the most appropriate choice in this situation.An ever-present consequence of using highly coherent illumination is speckle, arising from self-interference of the beam as it propagates through the optical system. This intensity variation is particularly problematic for lithography, since it directly affects the fidelity of pattern reconstruction. In SLM-based holography, the speckle noise is exacerbated by the digital noise arising from CGH calculation, as well as the discrete phase steps of the SLM pixels.

In SLM-based holography, the speckle noise is exacerbated by the digital noise arising from CGH calculation, as well as the discrete phase steps of the SLM pixels. However, it is straightforward to calculate multiple versions of a CGH that produce the same intensity pattern, by seeding the G-S algorithm with an initially random phase. Each of the resulting CGHs will then have different uncorrelated noise (and thus, speckle), and exposing them in rapid succession allows for an averaging effect that significantly improves the noise performance (estimated from statistical considerations to be a factor √𝑁 for 𝑁 CGHs).

The Holoeye PLUTO SLM used in this work is capable of cycling through CGHs at approximately 15 Hz, and each projected pattern uses 𝑁 = 10 different CGHs for speckle/noise reduction (see Figure 2).Eliminating Undiffracted LightOne of the most significant obstacles to the flexible and robust control of the light field in diffraction-based systems is eliminating interference from undiffracted light. Some fraction of the incident laser power reflects from the SLM without being modulated by the LC layer, generating an unwanted bright spot at the center of the build volume (with the intensity pattern that’s the Fourier transform of the SLM rectangular aperture). Of the many approaches to mitigating and eliminating interference from this uncontrolled light [30]–[33], we concluded that incorporating a beam block at the Fourier plane of the hologram projection lens is the most robust implementation in this context.

This approach is not pattern-dependent, nor does it require exhaustive SLM calibration and aberration compensation; once the beam block is aligned, additional adjustments are rarely necessary.The shape of the beam block is designed to maximize its efficiency in rejecting undiffracted light, while allowing maximum transmission of the holographically-controlled light field. We fabricated custom glass flats with a gold metal film to block the spot, patterned in a “star” shape, matching the dimensions of the undiffracted spot (see inset, Fig. 1). Then, by adding a controlled phase curvature (Fresnel lens) to the CGH, which increases the effective focal length of the projection lens, the desired hologram itself is displaced away from the beam1187Figure 3: Single-exposure holographic fabrication of a 3D structure.

(a) A representation of the incident beam geometry, with 45° mirrors reflecting two beams (which are subregions of the projected holographic field) along orthogonal directions to the original beam. (b) Photo of the cuvette before filling with resin, with paper screens at two sides for visualization of the projected pattern. (c) MATLAB point cloud representations of the propagating intensities for the three beams, and the region where they overlap in space. (d) The resulting part, after a single exposure of 12 s.block plane, thereby allowing use of the full hologram space, with virtually no interference from undiffracted light.Build Volume and PhotopolymerSince we are interested in 3D structure formation within the bulk of a liquid photopolymer, rather than in layers at a surface, this requires a resin formulation that is minimally absorptive.

The resin used here is poly(ethylene glycol) diacrylate (PEGDA, MW=250) with 0.04% (w/w) Irgacure 784, which is a titanocene free-radical photoinitiator with its absorbance spectrum extending to the 532 nm wavelength of the laser source. At the same time, the use of intersecting patterned beams requires optical access from at least three sides of the build volume. Therefore, to hold the resin volume of approx. 1 mL, we use a fluorometer cell (Starna Cells, 3-G-10) with 10×10×45 mm internal volume, and 1.25 mm thick polished glass sides and bottom. As with many resin-based systems, exposure doses vary depending on optical power density and geometry.

As with many resin-based systems, exposure doses vary depending on optical power density and geometry. With the 3-beam configuration, polymer parts were successfully produced using 12 s exposures at an estimated power of 30 mW/cm2 incident onto the cuvette from each side, corresponding to an estimated volumetric cure dose of ~45 mJ/mm3.Results and DiscussionFigure 3 summarizes the fabrication of a 3D cube structure using this holographic platform. Figure 4 on the next page shows several other fabricated geometries. We have chosen to fabricate millimeter-scale structures at this stage for simplicity of handling and geometric accessibility, but the part size and corresponding minimal feature scale can be readily adjusted by changing the magnification of the final projection optics. We intend to explore the resolution limits of this platform in future work.1188Figure 4: 3D structures fabricated with two-beam (a and b) and three-beam (c and d) exposures.

We intend to explore the resolution limits of this platform in future work.1188Figure 4: 3D structures fabricated with two-beam (a and b) and three-beam (c and d) exposures. (a) Shows the part in situ supported by the wall of the cuvette, however other parts form in the bulk volume, with no supporting structure. (b) This structure demonstrates the capability to produce structures with smoothly curving geometry without requiring layer-by-layer fabrication. Scale bars in all images are 2 mm.As evident from Figures 3 and 4, the three intersecting light patterns behave much like collimated beams, though this is not technically accurate. Rather, their depth of focus (DOF) is large enough such that there is negligible variation in the pattern over its 10 mm propagation through the build volume. The intersection volumes of the three patterns are cured in the resulting 3D structure.

The intersection volumes of the three patterns are cured in the resulting 3D structure. Thus, in the present implementation, we are constrained to “extruded” geometries whose cross-section along any one axis does not vary with single-exposure structures. Several approaches for moving beyond this limitation are noted below.The large DOF is a consequence of the SLM diffraction characteristics, as well as the final projection optics. These optics may be chosen with shorter focal lengths to reduce the effective system magnification, producing smaller-sized parts, and simultaneously increasing the numerical aperture (NA) at the build volume. This will improve axial resolution. However, this is fundamentally limited by the maximum diffraction angle available from the SLM pixels, as described by the diffraction grating equation 𝑎(sin 𝜃𝑑 − sin 𝜃𝑖) = 𝑛𝜆 where a is the grating pitch, n is the diffracted order, and 𝜃𝑖 and 𝜃𝑑 are the incident and diffracted beam angle, respectively.

For an SLM with 8 µm pixels diffracting 532 nm light, the angular spacing between diffracted orders is approximately 3.8°, corresponding to NA0.07 in air.Whyte and Courtial [34] provide a useful analysis of how NA limitations affect the accessible spatial frequencies of a holographically reconstructed geometry. The result, in brief, is that the spatial frequency (i.e. intensity variation) of the light field along the direction of propagation (𝑘𝑧) is necessarily very small for low-NA systems. We estimate that using this SLM, the maximum axial resolution is 0.25 mm at the HP and 0.5 mm at the cuvette due to the image magnification from lens pair L5 and L7 (see Fig. 1). However, this axial resolution is only attainable when lateral spatial frequency components in the pattern (𝑘𝑥 and 𝑘𝑦) are small, 2 where since these wave vectors in monochromatic light are related by 𝑘𝑥 𝑘0 = 2𝜋/𝜆.

Since the target intensity pattern is currently specified in a single plane, and the G-S algorithm is implemented for 2D fields, only the lateral spatial frequency components 𝑘𝑥 and 𝑘𝑦 are explicitly included in the calculation, and are privileged over any optimization of 𝑘𝑧.2 + 𝑘𝑦2 + 𝑘𝑧2 = 𝑘0The advantage of the approach described here is that the low spatial frequency in z can be compensated for with the higher spatial frequencies in x and y, by the folding of the light field, provided by the 45° mirrors. Taking advantage of this feature has already enabled the fabrication1189of the structures shown here. Moreover, we expect that the “extruded” geometry limitation will be overcome, by incorporating 3D Gerchberg-Saxton calculations of the phase field, and building the “folding” capability into the iterative algorithm.

This will allow the G-S algorithm to directly iterate on the 3D light field, providing further control on the image, and making more complex geometries routinely accessible with full control over all three spatial frequencies within the build volume.Conclusion and OutlookThis work constitutes the first demonstration of direct fabrication of unsupported aperiodic 3D structures, using a “volume-at-once” unit operation for photopolymer curing, without requiring a layer-by-layer process. We anticipate that this paradigm in 3D fabrication will bring considerable advances in fabrication speed and geometric flexibility to the field of photopolymer-based AM.

We anticipate that this paradigm in 3D fabrication will bring considerable advances in fabrication speed and geometric flexibility to the field of photopolymer-based AM. Additional geometric flexibility, as well as a larger overall build volume (or greater resolution in a small build volume), may be attained by using multiple SLMs illuminated by the same laser source, which project holographically interfering beams into the same resin volume.Although the part quality achieved thus far is low, this is not an inherent limitation for this method. Additional speckle noise reduction by de-speckling techniques, in addition to optimization of curing parameters with adjustment of the intensity distribution in the three sub- images will enable high-quality part fabrication.AcknowledgmentsThis work was performed under the auspices of the U.S.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceADDITIVE FABRICATION OF POLYMER-CERAMIC COMPOSITE FOR BONE TISSUE ENGINEERINGS. Ramesh*, M. Eldakroury*, I.V. Rivero*, M.C. Frank*Industrial and Manufacturing Systems Engineering, Iowa State University, Ames, IA 50011AbstractThe objective of this study is to manufacture chitosan-based biocomposite 3-D scaffolds through additive fabrication for promoting the regeneration of bone defects. Additive manufacturing has enabled the production of effective scaffolds by overcoming traditional limitations such as suboptimal distribution of cells, and poor control over scaffold architecture.

Additive manufacturing has enabled the production of effective scaffolds by overcoming traditional limitations such as suboptimal distribution of cells, and poor control over scaffold architecture. In this study, cryomilled biocomposites comprising of poly (lactic) acid (PLA), chitosan (CS) and tricalcium phosphate (TCP) provided the basis for the generation of hydrogels, which were then utilized for the fabrication of scaffolds with orthogonal (0, 90) geometry. Rheological studies were conducted using a rotational rheometer to identify the ideal hydrogel concentration for the continuous production of scaffolds. The scaffolds were fabricated using a 3-axis computerized numerical control (CNC) which was modified to function as a customized bioplotter. Scanning electron microscopy (SEM) was used to observe the morphology of the bioplotted scaffolds.

Scanning electron microscopy (SEM) was used to observe the morphology of the bioplotted scaffolds. Finally, a short-term stability (14 days) study was conducted to analyze the in vitro degradation behavior of the scaffolds in phosphate buffer saline (PBS).Keywords: Bioplotting, Chitosan, Cryomilling, Additive Fabrication, Bone Tissue Engineering1. IntroductionBone grafting is a surgical procedure that is frequently implemented to improve bone regeneration in a broad array of orthopedic procedures. Even though autografts are still considered the 'gold standard' for bone-grafting, extracting autologous bone tissue demands supplemental surgery and comes with limitations such as patient inconvenience, insufficient supply, and excessive costs [1]. Thus, conquering these limitations is essential to the development of efficient regenerative techniques to treat bone defects.

Thus, conquering these limitations is essential to the development of efficient regenerative techniques to treat bone defects. Over the last few decades, the scaffold-based therapy has emerged as an interesting alternative to overcome challenges associated with traditional tissue transplantation [2]. Natural and synthetic biodegradable polymers have been individually, or in a blended form, for the fabrication of scaffolds, the supportive structure that induces the regeneration of bone tissue [3]. Nevertheless, there is always an opportunity to explore new material combinations in an attempt to fabricate better scaffolds.Apart from desirable material properties, the design and fabrication of scaffolds have an impact on the success of the treatment. Traditional fabrication techniques such as solvent casting, fiber bonding, and freeze drying provide good macroscale features but often fail to match the complexity of natural human tissue [4].

Traditional fabrication techniques such as solvent casting, fiber bonding, and freeze drying provide good macroscale features but often fail to match the complexity of natural human tissue [4]. This necessitates the use of manufacturing techniques that would enable the production of scaffolds satisfying patient-specific needs.In an attempt to address the above-mentioned aspects, this study presents a preliminary feasibility report on the additive fabrication of 3D scaffolds using a novel biodegradable composite2365composed of poly (lactic acid) (PLA), chitosan (CS) and tricalcium phosphate (TCP). PLA, a synthetic polyester is known for its biocompatibility and suitable mechanical properties, however, the release of acid by-products during its hydrolytic degradation in vivo may have undesirable effects [5]. Likewise, CS, a naturally occurring polysaccharide, also possesses both desirable and undesirable qualities.

Likewise, CS, a naturally occurring polysaccharide, also possesses both desirable and undesirable qualities. It is used in the fabrication of biomedical scaffolds by virtue of its antimicrobial and biodegradable properties. However, its inappropriate mechanical and degradation properties have limited its exclusive use for medical applications [6]. The benefit of including TCP into a polymer matrix is twofold (a) it enhances material strength (b) also ensures the maintenance of a suitable ionic environment by neutralizing the acidic degradation of polyesters [5].

Taking together all these observations, it was hypothesized that the fabrication of biocomposite scaffold comprised of PLA, CS and TCP would be an effective solution to conceal the inherent drawbacks of these materials.In this study, additive fabrication was utilized for the fabrication of biocomposite hydrogel scaffolds due to its ability to produce intricate geometries which could be beneficial for cellular proliferation [7]. Nevertheless, the rheological behavior of the developed hydrogel needed to be evaluated as it can affect the printability and subsequently the stability of the scaffold [8]. Furthermore, the width of single layer filaments was used as a measurable parameter to quantify the suitability of the hydrogel for scaffold generation.In this manner, this study presents the fabrication approach of a novel powdered biocomposite hydrogel and serves as a feasibility report for the fabrication of complex 3D scaffolds for bone tissue engineering.2.

Materials and Methods2.1. Preparation of CS-based biocompositePLA (Purasorb PL 10; Purac, The Netherlands), CS, and TCP (Sigma-Aldrich, MO, USA) were cryomilled to generate powdered biocomposite. Cryogenic grinding was carried out using a freezer mill (SPEX, NJ, USA). Briefly, all the materials were weighed and transferred into a vial which was maintained at -196 using liquid nitrogen during the grinding process. The 20-minute grinding operation was divided into 4 cycles with a dwell time of 1 min between successive cycles. The composition of the generated biocomposite is as follows: 68%CS, 30%PLA and 2%TCP. The cryomilled biocomposite powders were subsequently used for the preparation of hydrogels.℃2.2. Characterization of cryomilled biocompositeX-ray diffraction (XRD) was used to confirm the occurrence of a homogenous blend of materials radiation source was used with the voltage and current adjusted to 30 kV and 15 mApost-cryomilling.

Rigaku Miniflex 600 XRD analysis unit (TYO, Japan) with a Cu-K ( respectively. 𝜆𝜆 Thermal characterization of the biocomposite powders was performed utilizing differential scanning calorimetry (DSC) (NETZSCH Instruments, MA, USA). Briefly, 10 mg of the sample followed by an was quenched to 10 /min to -30 for 1 min. Another isothermal step was performed for 3 mins after cooling isothermal step at 190 ℃ . The the samples down to -30 ℃= 0.154 nm)𝛼𝛼. The sample was then heated to 190℃℃. In the second heating cycle, the samples were heated to 250℃2366℃glass transition temperature and melting points were analyzed using the second heating cycle to facilitate direct comparisons of the samples.2.3. Rheological analysis of hydrogelsThree different concentrations (2.5%, 4%, and 5% w/v) of the biocomposite hydrogel were prepared using diluted acetic acid (2% v/v) (Sigma-Aldrich, MO, USA).

Rheological analysis of hydrogelsThree different concentrations (2.5%, 4%, and 5% w/v) of the biocomposite hydrogel were prepared using diluted acetic acid (2% v/v) (Sigma-Aldrich, MO, USA). Rheological measurements were obtained using an ARES-G2 rotational rheometer (TA Instruments, DE, USA).Two different routines were performed to fully understand the rheological behavior of the hydrogels. The first test was a steady shear sweep test intended to study the influence of shear rate ( ) of hydrogels. Curve fitting was then used to compute the consistency index, m (Pa.sn) and power law index, n (dimensionless). 𝛾𝛾̇ The second test was a 3-phase thixotropic test with the hydrogels being subjected to the application of shear rate periodically. The first phase lasted for 30 s with a shear rate of 0.1 s-1 (low), the second for 30 s with a shear rate of 200 s-1 (high) and the third for 60 s with a low shear rate again.

The first phase lasted for 30 s with a shear rate of 0.1 s-1 (low), the second for 30 s with a shear rate of 200 s-1 (high) and the third for 60 s with a low shear rate again. The change in viscosity was plotted as a function of time and the viscosity recovery behavior of the hydrogels were analyzed. The second study was designed to resemble the actual printing process which involves the periodic application of low and high shear rates.) on viscosity (𝜂𝜂2.4. Single layer testsThe ideal concentration for printing the biocomposite hydrogels was identified by printing and measuring single layered scaffolds. An air pressure of 30-40 psi provided uniform extrusion of all the concentrations. The width of the extruded filaments was obtained using a microscope (Leica, IL, USA).2.5.

An air pressure of 30-40 psi provided uniform extrusion of all the concentrations. The width of the extruded filaments was obtained using a microscope (Leica, IL, USA).2.5. Scaffold fabricationOrthogonal scaffolds (0, 90) were fabricated using a 3-axis computerized numerical control (CNC) machine (HAAS Automation, CA, USA), which was modified with the addition of a syringe head, pneumatic gauge and a controller to function as a bioplotter (Fig. 1a). The 2D sketches of the scaffolds and toolpaths were generated using Mastercam (CNC Software, CT, 40 psi was applied to ensure continuous deposition while a feed USA) (Fig. 1b). A pressure of rate of .40 inches/min was maintained constant. The scaffolds were fabricated using an 18 G needle with a distance of .3 inches from the petri-dish.

1b). A pressure of rate of .40 inches/min was maintained constant. The scaffolds were fabricated using an 18 G needle with a distance of .3 inches from the petri-dish. The hydrogels were dispensed into a medium containing sodium hydroxide (NaOH) (1 v/v%) (Sigma-Aldrich, MO, USA) and 100% high grade ethanol (Fischer Scientific, PIT, USA) in a 7:3 ratio.~2367Figure 1. (a) Bioplotting set-up created by modifying a 3-axis CNC machine, (b) Schematic showing the fabricated PLA+CS+TCP filaments in 0-90 layers with the associated toolpath generated using Mastercam.2.6. Scanning electron microscopyMorphology of the scaffolds were studied using scanning electron microscopy (SEM) (JCM- 6000Plus Neoscope JEOL, MA, USA). The pore and strand diameters were measured using JCM- 600 software version 1.1. The mean diameters were calculated using thirty measurements from three independently fabricated scaffolds and represented as meanSD.2.7.

The mean diameters were calculated using thirty measurements from three independently fabricated scaffolds and represented as meanSD.2.7. Short term stability analysis±In-vitro hydrolytic degradation was performed in a phosphate buffered solution (PBS) (pH=7.4) over a 14-day period. Scaffolds of 10mm x 10mm (4 layers) were immersed into a vial containing . The pH values were noted down every 15 ml of PBS and maintained in an incubator at 37 48 h using a pH meter (Sartorius, NY, USA).℃3. Results and Discussion3.1 X-ray diffractionThe ability of cryomilling to generate homogeneous blends of materials was confirmed using XRD. Figure 2 displays the diffractograms recorded for the as-received materials. PLA showed two distinct characteristic peaks, a strong diffraction peak at 2 and relatively weak peak . Apart from these, PLA largely exhibited the presence of amorphous regions at 2 and confirming the semi-crystalline nature of the PLA used in this study.

Apart from these, PLA largely exhibited the presence of amorphous regions at 2 and confirming the semi-crystalline nature of the PLA used in this study. The peaks at 16.38 𝜃𝜃 corresponded to d-spacing values of .54 nm and .47 nm which was consistent with previous 18.72 studies [9]. CS showed a sharp peak at 2 which related to d-spacing values of 0.45 nm and 0.95 nm respectively which was consistent with ° previously reported results [10]. TCP used in the study showed strong diffraction peaks at 2 which corresponded to d-spacing values of 0.34 nm, 0.28 nm values of 25.66 and 32.68 𝜃𝜃 and 0.27 nm. The peaks at 31.58 correspond to (0210) and (214) respectively [11].

The peaks at 31.58 correspond to (0210) and (214) respectively [11]. and 27.92 ° ° The diffractograms of the generated biocomposite displayed both the crystalline peaks of PLA ) peaks indicating the occurrence (16.38 of a homogeneous blend of the materials mixed.=16.38=18.72°𝜃𝜃°and a relatively broad peak at 2=19.44=9.34°𝜃𝜃𝜃𝜃, 31.58°° ) and TCP (31.58°) in addition to CS (19.44and 18.72°°°°2368°°Figure 2. XRD diffractograms of the materials used in this study.3.2. Differential scanning calorimetryDSC was used to identify the glass transition and melting temperatures of PLA, CS and the fabricated biocomposite. The thermograms from the second heating scan are shown in Figure 3. PLA showed an average glass transition temperature of 59°C with a melting point of 178°C and an endothermic cold crystallization peak at 164°C.

PLA showed an average glass transition temperature of 59°C with a melting point of 178°C and an endothermic cold crystallization peak at 164°C. The melting point of CS was not observed as it is a polymer of variable length and does not display a unique melting point. The thermogram of the biocomposite showed a melting temperature (177°C) similar to that of pure PLA which indicated the formation of similar lamellae population during cold crystallization [12].Figure 3. DSC second heating curves for PLA, CS and the cryomilled biocomposite.23693.3. Single layer testsThe single layer tests were conducted to measure the filament width as a function of material concentration. Figure 4a shows strands that were printed using 5% w/v biocomposite hydrogel and .15 mm which was greater than the diameter of the the mean strand diameter measured was 1.81 needle indicating that the hydrogel was experiencing die swell.

The filament width seemed to increase with decreasing hydrogel concentration as the 4% w/v hydrogel exhibited a mean value of 1.94 .15 mm. Moreover, the 2% w/v biocomposite was unable to form individual strands due to low viscosity (Figure 4b). The results are in agreement with previously reported studies where the hydrogel with a higher concentration maintained structural integrity in comparison with the lower concentration [8].±±Figure 4. (a) Single layer test performed with 5% w/v composite hydrogel, (b) 2.5% w/v hydrogel displaying inability to print individual strands.3.4. Rheological analysisAs previously mentioned, the steady shear sweep test was performed to analyze the effect of increasing shear rate on viscosity (Figure 5). The 2.5% w/v hydrogel showed a viscosity value of 5.67 Pa.s at 1 s-1 which dropped to 1.70 Pa.s at 100 s-1. The drop in viscosity for the 4% w/v concentration was more drastic where it dropped from 30.24 Pa.s at 1 s-1 to 3.51 Pa.s at 100 s-1.

The drop in viscosity for the 4% w/v concentration was more drastic where it dropped from 30.24 Pa.s at 1 s-1 to 3.51 Pa.s at 100 s-1. For the 5% w/v hydrogel, the viscosity dropped from 75.5 Pa.s to 2.7 Pa.s with increasing shear rate. All hydrogels exhibited typical shear thinning behavior with viscosities dropping with increasing shear rates. Based on the power law model, the constants ‘m’ and ‘n’ were obtained via curve fitting (Table 1). For a typical shear thinning liquid, the value of ‘n’ is expected to be smaller than 1 [13]. The values obtained from curve fitting clearly indicate that the higher concentration hydrogels exhibited a stronger shear thinning behavior.2370Figure 5. Viscosity of composite hydrogels as a function of shear rate at room temperature.Table 1.

Viscosity of composite hydrogels as a function of shear rate at room temperature.Table 1. ‘m’ and ‘n’ values obtained via curve fitting.Composition (w/v) 2.5% 4% 5%‘m’ (Pa.sn) 6.50 41.85 87.81‘n’ .74 .50 .28From a fabrication perspective, an ideal hydrogel needs to exhibit thixotropic behavior, lowering and recovering viscosities with the application and removal of shear rates. The recovery behavior of the synthesized hydrogels was analyzed using a 3-phase thixotropic test as described in section 2.3. As seen in figure 6, all the hydrogels studied exhibited considerable recovery of viscosity with the removal of shear rate. The fluctuations in the viscosity values of the 5% w/v hydrogel were explained by the reduction in the average distance between particles and an increase in particle-particle interaction and hydrodynamic effects [14]. The 5% w/v biocomposite hydrogel exhibited a slightly lower percentage of recovery (79%) when compared to the 4% w/v (86%).

The 5% w/v biocomposite hydrogel exhibited a slightly lower percentage of recovery (79%) when compared to the 4% w/v (86%). However, the evaluation of the aggregate results derived from filament width and rheological characteristics justify the selection of the 5% w/v biocomposite for the fabrication of scaffolds.2371Figure 6. Viscosity as a function of time at constant shear rate for composite hydrogels.3.5. Visual characterizationSEM was used to study the morphology of bioplotted orthogonal scaffolds. Figure 7a shows a micrograph of the fabricated biocomposite scaffold. The printed scaffolds exhibited a strand diameter of 318 m. It was hypothesized that the presence of PLA in the scaffold would counter the effects of scaffold swelling caused by the hydrophilic chitosan. The scaffold exhibited a slightly rough texture which was seen as a positive factor that could encourage cell attachment [15].

The scaffold exhibited a slightly rough texture which was seen as a positive factor that could encourage cell attachment [15]. Figure 7b shows the scaffold fabricated with 5% w/v biocomposite hydrogel.30.15m and a pore diameter of 113112±𝜇𝜇±𝜇𝜇Figure 7. (a) Micrograph of the bioplotted biocomposite scaffolds (4 layers), (b) Scaffold fabricated from 5% w/v biocomposite hydrogel.23723.6. Short-term stability analysisThe short-term stability was conducted with the intention of determining the structural integrity and the degradation behavior of the scaffolds in a biological environment. To establish a control for comparison, pure chitosan scaffolds were fabricated and subjected to testing under the same conditions. The results from the pH analysis, as seen in Figure 8 indicated that the biocomposite scaffolds maintained a neutral pH while the pure chitosan scaffolds exhibited a slightly alkaline pH.

The results from the pH analysis, as seen in Figure 8 indicated that the biocomposite scaffolds maintained a neutral pH while the pure chitosan scaffolds exhibited a slightly alkaline pH. It was noted that the addition of PLA can be tailored to adjust the pH of the degradation process to make it patient-friendly. In other words, the acidic degradation of PLA was able to counter the alkaline degradation of chitosan to ensure the existence of a stable neutral pH [16].Figure 8. pH of PBS as a function of time for the fabricated scaffolds.4. ConclusionsThis study investigated the preparation and characterization of a novel biocomposite hydrogel for the fabrication of bone tissue engineering scaffolds. The rheological and filament width characterization confirmed that 5% w/v biocomposite hydrogel was the most suitable from the lot evaluated. Also, the SEM observations confirmed the production of a controlled scaffold structure.

DITIVE MANUFACTURE OF LARGE STRUCTURES: ROBOTIC OR CNC SYSTEMS?Yashwanth K. Bandari*, Stewart W. Williams, Jialuo Ding, and Filomeno MartinaWelding Engineering and Laser Processing Centre (WELPC), Cranfield University, Cranfield MK43 0AL, Bedford, United Kingdom *y.k.bandari@cranfield.ac.ukAbstractAdditive manufacture of metre scale parts requires direct feed processes such as blown powder or wire feed combined with lasers or arcs. The overall system can be configured using either a robotic or Computer Numerical Controlled (CNC) gantry system. There are many factors that determine which of these is best and this will be presented in this paper. Some factors are inherent to the specific process type such as accuracy/resolution and any requirement for reorientation of the feedstock and heat source. Other factors depend on the particular application including material type, shielding options, part size/complexity, required build strategies and management of distortion.

Other factors depend on the particular application including material type, shielding options, part size/complexity, required build strategies and management of distortion. Further considerations include the incorporation of ancillary processes such as cold work, machining or inspection. The relative influence of these factors will be discussed. Cost implications for the different approaches will be highlighted based upon the type of process being utilized. Examples are provided where both robotic and CNC options have been evaluated and the best solution found.Keywords: Additive Manufacture, Direct feed, Robotics, Cost1. IntroductionAdditive Manufacturing (AM) has over 100 years of history - the first patent was filed in 1920 by Baker [1]. This technology is becoming an increasingly important in the manufacturing field for direct fabrication of structural components.

This technology is becoming an increasingly important in the manufacturing field for direct fabrication of structural components. The basic and common principle of these AM technologies involves the direct production of 3D objects from computer aided design (CAD) by adding materials in layers. Nowadays, modern manufacturing industries, such as automotive and aerospace, are continuously seeking applications and systems that enable direct production of large and relatively complex full density metal parts, with low production volumes, high structural integrity, accurate dimensions, and that can be directly used in operational systems.AM is also known as rapid manufacturing [2] or rapid prototyping [3]. Unlike conventional manufacturing techniques such as machining and stamping that fabricate products by removing materials from a larger stock or sheet metal, AM creates the final shape by adding materials.

It has the ability to make efficient use of raw materials and produce minimal waste while reaching satisfactory geometric accuracy. Using AM, a design in the form of a computerized 3D solid model can be directly transformed to a finished product without the use of additional fixtures and cutting tools. This opens up the possibility of producing parts with complex geometry that are difficult to obtain using material removal processes. Lastly, AM is a cost-competitive approach for fabricating components made of expensive materials such as titanium and nickel alloys in aerospace industries, where such components suffer an extremely low fly-to-buy ratio [4].AM technologies are mainly classified into powder bead fusion, directed energy deposition, binder jetting and sheet lamination as provided in Table 1. Typical additive materials are metal powder and metal wire.

Typical additive materials are metal powder and metal wire. With regard to how the additive material is supplied, currently popular AM technologies can be classified as either a powder-feed/-bed process or a wire-feed process The majorities of the research in AM have been focused on the powder-feed/-bed AM, where the laser or electron beam equipment is usually used as the power source. Powder-based AM technologies are mainly classified into powder bead fusion, directed energydeposition, binder jetting and sheet lamination as provided in Table 1. Typical additive materials are metal powder and metal wire. With regard to how the additive material is supplied, currently popular AM technologies can be classified as either a powder-feed/-bed process or a wire-feed process. The majorities of the research in AM have been focused on the powder-feed/-bed AM, where the laser or electron beam equipment is usually used as the power source.

The majorities of the research in AM have been focused on the powder-feed/-bed AM, where the laser or electron beam equipment is usually used as the power source. Powder-based AM techniques generally involve a complex non-equilibrium physical and chemical metallurgical process, which exhibits multiple modes of heat and mass transfer, and in some instances, chemical reactions. A comprehensive review on the materials design, process control, property characterisation and metallurgical theories for laser sintering, laser melting and laser metal deposition of a wide variety of metallic powders has been reported recently [4]. Table 2 provides a comparison of the basic features between powder-feed/-bed and wire-feed process.Table 1 Classification of Additive Manufacturing [4]ClassificationTerminologiesRef.

Table 2 provides a comparison of the basic features between powder-feed/-bed and wire-feed process.Table 1 Classification of Additive Manufacturing [4]ClassificationTerminologiesRef. MaterialPowder bed fusionDirect metal laser sintering (DMLS) Electron beam melting (EBM) Selective laser sintering (SLS) Selective laser melting (SLM)Metal powderDirected energy depositionElectron beam freeform fabrication (EBF3 ) Laser engineered net shaping (LENS) Laser consolidation (LC) Directed light fabrication (DLF) Wire and arc additive manufacturing (WAAM)Metal powder, metal wireBinder jettingPowder bed and inkjet 3D printing (3DP)Metal powderSheet laminationLaminated object manufacturing (LOM) Ultrasonic consolidation (UC)Metal laminate metal foilThe powder-feed/-bed approach is better developed due to its capability of fabricating parts with high geometrical accuracy.

The typical layer thickness in powder-feed/-bed technology is 20–100 µm, and the completed components can achieve a dimensional accuracy of ±0.05 mm and surface roughness of 9–16 µm. In addition, it is possible to produce parts with functionally graded materials (FGM). However, the deposition rate of the powder-feed/-bed technology is extremely low, typically around 10 g/min, which limits its application in fabricating median to large-sized components. In wire-feed AM, a metal wire is used as supply material instead of metal powder. Depending on the energy source used for metal deposition, wire-feed AM can be classified into three groups, namely: laser-based, arc welding-based and electron beam-based . Wire-feed AM has higher material usage efficiency with up to 100 % of the wire material deposited into the component. Therefore, it is a more environmental friendly process, which does not expose operators to the hazardous powder environment.

Therefore, it is a more environmental friendly process, which does not expose operators to the hazardous powder environment. Compared with the powder-feed process, it has a much higher deposition rate of up to 2500 cm3/h (330 g/min for stainless steel)[4]. It reveals that there is a trade-off between high deposition rate and high resolution while selecting which type of AM process to use for a certain component.Table 2 Comparisons of some AM processes [4]Additive materialsPowderProcessLC SLM SLS DLFLayer thickness (µm) Deposition rateN/A 20-100 75 200(kg/hr) 0.06-18 N/A ~0.006 0.06Dimensional accuracy (mm) ±0.025–±0.069 ±0.04 ±0.05 ±0.13Surface roughness (µm) 1-2 9-10 14-16 ~20WireWAAM EBF~1500 N/A0.72- 4 Up to 19±0.2 Low200 High2. Means of fabricating AM partsAM parts require direct feeding of blown powder or wire feeding combined with laser or arcs.

Means of fabricating AM partsAM parts require direct feeding of blown powder or wire feeding combined with laser or arcs. They are generally manufactured using industrial robots or on Computer Numerical Control (CNC) gantry systems. Recent developments in AM reflect the requirements for flexibility to adapt to changes taking place in the market, in the society and in the global economic environment. Major objectives for producing AM parts today are evolving. Users’ requirements indicate that emphasis should be given to low-volume and large variety, even in high volume production industry, to face global competition. Flexibility is also required to use the same facility for the minor or major model changes that come in effective life span of the equipment [5].The synergic integration of deposition unit with the CNC machine is independent of its age and make.

Generally, the integration is done in such a way that the deposition acts an additional feature without disturbing the other capabilities of CNC machine. The other issues which are generally sorted are mounting of the torch on the side of the spindle head so that weld deposition is controlled through the same CNC controller, ability to retract the torch during starting and end of the deposition, appropriate safeguards to protect the machine elements from occasional spatter. The retraction of the torch is demonstrated in the figure 1. The full set up of the CNC gantry system with the deposition unit is shown in the figure 2.Figure 1: Retracting the torch between the limitsFigure 2: CNC gantry system with deposition power sourceIn case of robotic systems, the torch is attached to the wrist of the robot which has two or three axes of motion. The robot is programmed to move the torch along the path in a given orientation.

The robot is programmed to move the torch along the path in a given orientation. The majority of the industrial robots are actuated by linear, pneumatic, or hydraulic actuators, and/or electric motors.Table 3 shows the list of some universities/institutes where either of robotic system or CNC gantrysystem is used for deposition (not limited to), along with the deposition methodology.Table 3 List of universities/institutes either robotic or CNC system is usedCNC gantry systemRobotic systemLos Alamos laboratory (DLF) Southern Methodist University (DMD) Korean Institute of Science and Technology Indian Institute of Technology Bombay (WAAM) Wuhan University (Hybrid Plasma Dep & milling)University of Nottingham (3D welding) Cranfield University (WAAM) TWI (DMD) Renishaw (DMLS) Fronius, Lockheed Martin (WAAM)There are many factors which determine which one is best and this will be presented in the paper.

Some factors are inherent to the specific process type such as accuracy/resolution and any requirement for reorientation of the feedstock. Other factors depend on the particular application including material type, shielding options, part size, required build strategies and rotation of the part. Further considerations include the incorporation of ancillary processes such as cold work, machining or inspection. The relative influence of these factors will be discussed. Cost implications for the different approaches will be highlighted based upon the type of process being utilized. Examples are provided where both robotic and CNC options have been evaluated and the best solution found.3. Factors influencing which is besta) Accuracy: An industrial robot can fulfill the needs of todays and tomorrows industry for time and cost efficient, yet flexible mean of material processing.

Factors influencing which is besta) Accuracy: An industrial robot can fulfill the needs of todays and tomorrows industry for time and cost efficient, yet flexible mean of material processing. Most industrial robots are constructed as a cantilever, in which each of the arms is supported by motors, brakes and reduction gears, they struggle to achieve high positioning accuracy level, being limited to 0.5-2 mm [6] and at the same time are more prone to disturbances from the process forces. Thus, robot is not robust and accurate compared to the CNC system.An investigation was done to measure the speed of the robot whilst moving along the sharp corners. The robotic movement was captured by a high speed camera with 100 frames per second. The video was analyzed by splitting into frames, and the robotic position in each frame was plotted on a graph sheet and the distance between successive positions was calculated.

The video was analyzed by splitting into frames, and the robotic position in each frame was plotted on a graph sheet and the distance between successive positions was calculated. It was approximated to be a straight line between two successive positions because of large frame rate. If (x1, y1) and (x2, y2) represent 2 points in successive frames then the distance between them is given by the formula below:From the distance calculated, it was seen that at the corners, the distance between points in successive frames decreased, when compared to the distance between points when the robot was moving in a straight line path, implying that the speed of the robot decreased. It was also dependent on the robotic speed set. It was seen that the change in speed increased with the increase in robotic speed.

It was also dependent on the robotic speed set. It was seen that the change in speed increased with the increase in robotic speed. Figure 3 shows the graph between % change in travel speed of the robot at the corners and set travel speed of the robot.Figure 3 Graph between change in travel speed at corners and travel speed setb) Shielding Vs material type: Shielding gases are of considerable significance in the protection of molten pool from atmospheric contamination during the deposition process. These gases play an important role in a number of aspects. Shielding gases in arc and laser processes have a remarkable effect on the overall performance of the welding system. It is also dependent on the process type and the material type. For the processes like GMAW, local shielding is required and for Plasma Arc Welding (PAW) with titanium, global inert atmosphere is needed.

For the processes like GMAW, local shielding is required and for Plasma Arc Welding (PAW) with titanium, global inert atmosphere is needed. For the materials like mild steel, stainless steel, high strength steel, aluminum, Inconel, which are the commonly used ones; local shielding is required to avoid oxidation of molten pool, with varying combinations of Carbon-di-oxide and oxygen levels.For process like PAW and materials like titanium, for global inert atmosphere, different options are tents, enclosures and local shielding. It is comparatively easy to set up the enclosures with a robot, compared to on a CNC system, because of the space occupied by the CNC and cost for setting up. Figure 4 shows some of the examples of the enclosures with robot.Tent Enclosure Figure 4 Different enclosures to avoid oxidationc) Process type: Part building is mainly done using GMAW, TIG and PAW.

During the deposition process, wire feed orientation influences drop transfer and the quality of the deposit. In case of GMAW, wire is fed automatically through the torch (head), driving it through the conduit. Back feeding, side feeding, front feeding has all been tried in case of TIG and PAW. Thus, torch always needs to be rotated to guarantee the consistency of the bead geometry [7]. It is difficult to change the orientation of wire frequently in case of CNC system. However, with the robot, it is easy to rotate the torch, so that the wire can be fed in any orientation. Figure 5 shows the torch arrangement for a TIG set up.Figure 5 TIG torch with wire feederd) Deposition strategy: In order to obtain a geometrically stable deposited product, it is essential to deposit the metal using the strategies which results in a flat surface to be good base for the subsequent layer or nullify the amplification of distortion, residual stresses, dimensional inaccuracies, weld defects etc.

It is generally found that as the number layers deposited are more, the geometric accuracy and precision as well as the thermal conditions within the object deteriorate. They cause major problems because they decrease the level of quality and subsequently the path design gets complicated [8].Traditionally, wider walls are built by depositing metal next to each other in a parallel fashion, in a required number of passes. However, the arc needs to be turned off and on at the end of each pass, which results in higher fabrication time. Due to the considerable time between the adjacent passes, the metal becomes cold and thus, there is a high possibility of finding voids as defects. Therefore, to avoid these conditions, a new strategy proposed is called parallel linear square rounded strategy which would overcome all the drawbacks of parallel strategy.

Therefore, to avoid these conditions, a new strategy proposed is called parallel linear square rounded strategy which would overcome all the drawbacks of parallel strategy. Figure 6 shows the parallel and parallel linear square rounded pattern.Figure 6 Parallel pattern (left), linear parallel square rounded pattern (right)When walls are built using the parallel linear square rounded pattern on CNC and on robot, there was a difference in surface profile between the samples. Figure 7 shows the weld samples built using robot and on CNC, with same conditions.Figure 7 Samples deposited using robot (left) and on CNC system (right)From the deposited samples, it is clear that the surface profile is not the same. The one deposited on the CNC is more flat compared to the one built using robot. There is an extra material deposited at the corners when built using robot. This is due to the change of speed of robot at the corners. Travel speed of 7 mm/s was used to build the sample.

There is an extra material deposited at the corners when built using robot. This is due to the change of speed of robot at the corners. Travel speed of 7 mm/s was used to build the sample. Figure 3 shows that with 7 mm/sec travel speed, there is a decrease of 25 % in travel speed, which means there is a 25% extra material deposited. However, the movement of torch is governed by DC motors in case of CNC system and has a better control and thus, there is hardly any change in travel speed at the corners. Therefore, CNC is preferred to robotic system in this case.e) Part rotation: Part rotation is one of the essential feature of part building, if part needs to be built double sided or part having overhang features which require 4/5 axis kinematics, counter distortion/residual stresses. In case of CNC, it is not possible to rotate the large bed. Some simple 4/5 axis CNC systems are capable but the cost of them are pretty high.

In case of CNC, it is not possible to rotate the large bed. Some simple 4/5 axis CNC systems are capable but the cost of them are pretty high. However, with robot, part rotator can be incorporated, which would work independently or can have coordinate motion installed in it. Figure 8 demonstrates a U-shaped component where part rotation is necessary to build and also a part built double sided.Figure 8 Part rotation examplesf) Cost comparison: The cost of a simple robot starts from 40,000 £ and with additional features like rotator, machining, so on and so forth the cost of it increases to 160,000 £. However, the cost of the simple 3 axis CNC machine starts from 100,000 £ and with additional features it increases to nearly 1,500,000 £, which is nearly 10 times higher than the cost of the robot. Table shows the various costs involved in building a rib with titanium, taking into consideration the higher end robotic and CNC systems.

Table shows the various costs involved in building a rib with titanium, taking into consideration the higher end robotic and CNC systems. Thus, it can be concluded that cost of robotic system is less compared to cost of CNC gantry system.Table 4 Cost comparison between robotic system and CNC system for Titanium ribFeature Hardware Software Labor cost Hourly cost of the cell Cost of the part Specific cost (£/kg)Cost of robotic system (£) 0.16 M 90000 45 78 5043 249Cost of CNC gantry system (£) 1.5 M 90000 45 246 7382 364g) Additional Processes: Incorporating additional processes like rolling, machining, NDT inspection are comparatively easier on a CNC system rather than on a robot. The robotic systems are not that rigid enough to absorb forces and pressure arising from the above listed additional processes.

The robotic systems are not that rigid enough to absorb forces and pressure arising from the above listed additional processes. Figure 9 shows rolling process incorporated on a CNC machine.Figure 9 Rolling rig added as an additional feature to CNC systemh)Scale: It is possible to build large parts on CNC gantry systems. It all depends upon the dimensions of the bed. Figure 10 shows a 5 m length wall deposited on a CNC system.Figure 10 5 m long wall built on CNC gantry systemHowever, using robot as well, it is possible to build large structures using multiple robots or using robots on linear drives to facilitate the moving the robot as a whole to deposit at various places and increase its working envelope. Figure 11 shows multiple robots building an aluminum rib.Figure 11 Multiple robots building a long Aluminum rib4. ConclusionFrom the factors listed above, it is clear that usage of either CNC gantry system or robotic system is purely dependent upon application.

ConclusionFrom the factors listed above, it is clear that usage of either CNC gantry system or robotic system is purely dependent upon application. Robotic system is preferred when the part is to be built with lower cost, very long, involves rotating of the part, to be built in a global inert environment. On the other side, CNC gantry system is preferred when the part is to be built with high accuracy levels, involves additional processes like rolling, machining etc.References[1] Almeida, P. (2012), Process control and development in wire and arc additive manufacturing, PhD Thesis,Cranfield University. U. K[2] Levy, G., Schindel, R., Kruth, J.P., (2003), Rapid Manufacturing and Rapid Tooling with LayerManufacturing (LM) Technologies, State of The Art and Future Perspectives; CIRP Annals Vol. 52/2/2003; p.

Hallgren, L. Pejryd and J. Ekengren, "(Re)Design for Additive Manufacturing," in CIRP DesignConference, Stockholm, 2016.[2] S. Crump, "Apparatus and Method for CreatingThree-Dimensional Objects". United States of AmericaPatent US5121329A, 09 June 1992.[3] L. Gibson and M. Ashby, "Cellular Solids: Structure And Properties", Cambridge University Press, 1997.[4] T. Bhardwaj, S. Singh and M. Shukla, "Finite Element Modeling and Analysis of Implant Scaffolds," in International Conference on Advances in Mechanical, Industrial, Automation and Management Systems, Allahabad, 2017.[5]J. Hawreliak, J. Lind, B. Maddox, M. Barham, M. Messner, N. Barton, B. J. Jensen and M. Kumar, "Dynamic Behavior of Engineered Lattice Materials," Scientific Reports 6., 2016.[6] A. Panesar, M. Abdia, D. Hickman and I. Ashcroft, "Strategies for Functionally Graded Lattice Structures Derived Using Topology Optimisation For Additive Manufacturing," Additive Manufacturing, vol. 19, pp.

Abdia, D. Hickman and I. Ashcroft, "Strategies for Functionally Graded Lattice Structures Derived Using Topology Optimisation For Additive Manufacturing," Additive Manufacturing, vol. 19, pp. 81-94, 2018.2352[7] P. Terriault and V. Brailovski, "Modeling and Simulation of Large, Conformal, Porosity-graded and Lightweight Lattice Structures Made by Additive Manufacturing," Finite Elements in Analysis and Design, vol. 138, pp. 1-11, 2018.[8]J. Mutua, S. Nakata, T. Onda and Z. C. Chen, "Optimization of Selective Laser Melting Parameters and Influence of Post Heat Treatment on Microstructure and Mechanical Properties of Maraging Steel," Materials and Design, vol. 139, pp.

8(7), 2018.[10] B. Mooney, K. I. Kourousis, R. Raghavendra and D. Agius, "Process Phenomena Influencing the Tensile and Anisotropic Characteristics of Additively Manufactured Maraging Steel," Materials Science & Engineering A, vol. 745, pp. 115-125, 2019.[11] Vehicles with Regard to Door Latches and Door Retention Components. [Art]. Status of United NationsRegulation, 2010.[12] "Simulation Driven Design with Inspire," Altair University, 2018.2353.

DITIVE MANUFACTURING (AM) AND NANOTECHNOLOGY: PROMISES AND CHALLENGESOlga S. Ivanova1, Christopher B. Williams2, Thomas A. Campbell1*1.Institute for Critical Technology and Applied Science (ICTAS), Virginia Tech, Blacksburg, Virginia, 24061, USA2. Department of Mechanical Engineering & Department of Engineering Education, Virginia Tech, Blacksburg, Virginia, 24061, USACorresponding Author, tomca@vt.edu, 540-231-8359AbstractThe narrow choice of materials used in Additive Manufacturing (AM) remains a key limitation to more advanced systems. Nanomaterials offer the potential to advance AM materials through modification of their fundamental material properties. In this paper, the authors provide a review of available published literature in which nanostructures are incorporated into AM printing media as an attempt to improve the properties of the final printed part.

Specifically, we review the research in which metal, ceramic, and carbon nanomaterials have been incorporated into AM technologies such as stereolithography, laser sintering, fused filament fabrication, and three-dimensional printing. The purpose of this article is to summarize the research done to date, to highlight successes in the field, and to identify opportunities that the union of AM and nanotechnology could bring to science and technology.IntroductionAdditive Manufacturing (AM) is a group of emerging technologies that create objects from the bottom-up by adding material one cross-sectional layer at a time [1]. For the past three decades, researchers and engineers have focused on improving old and creating new techniques, as well as developing novel materials.AM methods have several advantages over traditional manufacturing techniques.

First, AM offers “design freedom” for engineers; because of its additive approach, it is possible to build geometries that cannot be fabricated by any other means. Moreover, it is possible with AM to create functional parts without the need for assembly. Also, AM offers reduced waste; minimal use of harmful chemicals, such as etching and cleaning solutions; and the possibility to use recyclable materials.Despite these design and environmental advantages, the adoption of AM as a means for fabricating end-use components has historically been dampened by the technologies’ narrow selection of available materials. The majority of materials used presently by modern AM techniques are proprietary polymers. There are also possibilities to use some metals (such as steel alloys and titanium) and ceramics.

The majority of materials used presently by modern AM techniques are proprietary polymers. There are also possibilities to use some metals (such as steel alloys and titanium) and ceramics. While there are different types of materials that can be used in AM, the material properties are typically not as strong as their conventionally manufactured counterparts due to the anisotropy caused by the layer-by-layer approach. Therefore, options within these materials genres are usually limited to applications of models for form/fit testing, functional testing, presentation models, prototypes and non-load bearing products. Furthermore, variations in build process parameters, and in some AM processes,ambient conditions, can result in variations between parts built on different machines of the same AM technology.

Furthermore, variations in build process parameters, and in some AM processes,ambient conditions, can result in variations between parts built on different machines of the same AM technology. Overcoming these issues will require advances in both process control approaches and material selection.The integration of nanotechnology with AM has the potential to both complement existing techniques and to create wholly new nanocomposites; it is thus a promising approach to alleviating some of the technologies’ limitations. “Nanotechnology is science, engineering, and technology conducted at the nanoscale, which is about 1 to 100 nanometers” [2]. When shrinking the size scale from macro to nano, or bulk to molecule, materials can change their fundamental properties. For example, at the nanoscale, objects can exhibit unique optical [3-6], thermal [7], and electrochemical [8-10] properties that differ from the properties of the bulk material or molecules.

For example, at the nanoscale, objects can exhibit unique optical [3-6], thermal [7], and electrochemical [8-10] properties that differ from the properties of the bulk material or molecules. These properties strongly depend on the size and the shape of nanostructures [11, 12]. There are also a wide variety of nanomaterials, including carbon nanotubes (CNTs), nanowires (NWs), buckyballs, graphene, metal nanoparticles (NPs), and quantum dots (QDs). All these materials possess unique characteristics that allow applications in areas such as sensing [13-15], separation [16], plasmonics [17-19], catalysis [20-25], nanoelectronics [26-28], therapeutics [29-31], and biological imaging and diagnostics [32-38]. Thus, the addition of nanomaterials to AM printing media could enable the creation of entirely new composites possessing unique properties and lead to expansion of application areas of AM. Published research combining AM and nanomaterials thus far is limited.

Published research combining AM and nanomaterials thus far is limited. Nevertheless, articles have shown that the introduction of inorganic nanostructures such as CNTs, metal NPs, and ceramics can significantly affect sintering characteristics and final mechanical properties of the printed parts. This literature review summarizes the work performed to date in the area of incorporating nanomaterials into AM and describes the effects of the addition of nanostructures on the physical properties of the final printed part. We conclude with a discussion of challenges and opportunities remaining in the union of AM and nanotechnology.Metal NanoparticlesMetal nanoparticles (NPs) are of great interest to researchers because their optical, thermal and electrochemical properties depend significantly on their size. Metal NPs can be synthesized with different shapes – spheres, cubes, rods – depending on the process.

Metal NPs can be synthesized with different shapes – spheres, cubes, rods – depending on the process. Figure 1-A shows the photographs of silver and gold nanoparticle solutions synthesized by chemical reduction method in the presence of citrate ions [39, 40]. Figure 1-B represents a typical SEM (Scanning Electron Microscopy) image of the spherical Ag nanoparticles (d=10-15 nm) deposited on an electrode surface. Metal NPs have found applications in sensing, catalysis, bio- imaging, drug delivery, and microelectronics.Recently metal NPs have been utilized for the purposes of improving the sintering characteristics and creation of new printing materials for AM. One of the first reports on using metal nanoparticles for AM is by Crane, et al., in 2006 [41]. Crane used iron NPs (diameter = 7 to 10nm) to improve the sintering quality of steel parts created by 3D printing (3DP). The test specimens were printed using traditional binder and a steel powder ranging from 63 to 90µm in diameter.

The test specimens were printed using traditional binder and a steel powder ranging from 63 to 90µm in diameter. Different approaches for infiltration of NPs were tested, such as multiple applications of nano-binders, the rate of drying, drying under a magnetic field, and the use of oxidized and unoxidized NPs. It was found that the addition of a nano-binder to the printed parts decreased deflection and shrinkage. Figure 2 shows the comparison of creep deflection of two printed partsFigure 1. (A) Photographs of Ag and Au nanoparticle solutions. (B) SEM image of the Ag NPs deposited on a Glass/ITO electrode surface.during sintering, the first with nano-iron treatment and the second without nano-iron treatment. The deflection of the part treated with nano-iron was reduced by 95% and shrinkage was reduced by 60%.

The deflection of the part treated with nano-iron was reduced by 95% and shrinkage was reduced by 60%. Crane suggests that the application of the nano-binder multiple times decreased the deflection of the final parts by enlarging the bonds between bigger particles and curing the cracks formed after the sintering process.Figure 2. Comparison of deflection of stainless steel bars with and without nano-iron treatment. Treated bar received four applications and was dried under a magnetic field. Ref. [41]Bai, et al., [42] added silver NPs (diameter = 30 nm) directly into a binder solution to create a silver NP suspension for 3DP. It was shown that the nanosilver suspension could be sintered at 300℃. Low sintering temperatures of metal nanoparticles are due to the thermodynamic size effect [7]. The quality of the parts printed with nanobinder and with traditional binder was compared.

Low sintering temperatures of metal nanoparticles are due to the thermodynamic size effect [7]. The quality of the parts printed with nanobinder and with traditional binder was compared. It was demonstrated that the addition of Ag NPs improved the sintering characteristics of the final part. Parts printed with nanobinder showed less distortion and shrinkage compared to a pure binder system.Use of highly concentrated Ag NPs suspension for fabrication of microelectrodes using Omnidirectional Printing was shown by Ahn, et al. [43]. Ag nanosuspensions with solid loadings of >70% were produced by concentrating the NPs with ethanol and centrifugation. These highly concentrated NP inks were then used for fabrication of Ag microelectrodes. It waspossible to use nozzles with a diameter as narrow as 1µm to print silver nanosuspensions (70% loading) without clogging.

It waspossible to use nozzles with a diameter as narrow as 1µm to print silver nanosuspensions (70% loading) without clogging. Printing of planar and flexible microelectrodes, and also multilayered and arched architectures, was achieved using these inks (Figure 3). Electrical resistivity and effects of mechanical bending on the electrical performance of fabricated microelectrodes were tested. Electrical resistivity of these electrodes decreased with an increase in annealing temperature and was close to the value for bulk silver when annealed at 250℃. Printed electrodes demonstrated steady electrical response for at least 750 bending cycles.Figure 3. (A) SEM images of planar arrays of silver microelectrodes patterned with 1- (left), 5- (center), and 10-mm (right) nozzles. (B) SEM images of multilayer silver microelectrodes patterned with 5- (top left), 10- (bottom left), and 30-mm (right) nozzles. (C) a silver interconnect arch printed over an electrode junction. Ref.

(B) SEM images of multilayer silver microelectrodes patterned with 5- (top left), 10- (bottom left), and 30-mm (right) nozzles. (C) a silver interconnect arch printed over an electrode junction. Ref. [43]These examples show that the addition of metal NPs can significantly affect sintering characteristics of the final parts, decreasing their shrinkage and distortion. Also, highly concentrated metal NP inks are promising materials for the fabrication of microelectrodes and microconnectors; they could find use in microelectronics, solar cells, batteries, etc.Carbon Nanotubes and NanofibersCarbon nanotubes (CNTs) are allotropes of carbon. The structure of CNTs is cylindrical; each carbon atom has three neighbors, as in graphite. CNTs are considered as rolled-up graphene sheets (graphene is a one-atom-thick sheet of graphite).

The structure of CNTs is cylindrical; each carbon atom has three neighbors, as in graphite. CNTs are considered as rolled-up graphene sheets (graphene is a one-atom-thick sheet of graphite). There are three primary CNT forms – armchair, zig-zag, and chiral – as shown in Figure 4; characterization of CNTs can be challenging and may require new instruments [44]. Carbon nanotubes are among the strongest and stiffest materials known [45, 46], and they possess unique electrical and thermal conductivities [47-50]. CNTs have been consequently used as additives to polymers to improve mechanical, thermal, and electrical properties [51, 52]. They have found applications in hydrogen storage, field emission devices, sensors, and scaffolds for damaged nerve regeneration,etc. [ 53-55]. Here we describe work performed with CNTs as additives to the materials for Stereolithography, Laser Sintering and Direct Writing.Figure 4.

[ 53-55]. Here we describe work performed with CNTs as additives to the materials for Stereolithography, Laser Sintering and Direct Writing.Figure 4. Single walled CNTs structures: (A) armchair, (B) zig-zag, and (C) chiral.Sandoval, et al. [56, 57] used multiwalled carbon nanotubes (MWCNTs) with an average diameter of 30nm and lengths in the range of 5 to 20µm as an additive to stereolithography (SL) resin. Results showed that the addition of MWCNTs increase tensile stress and fracture stress by ~5.7% and ~26%, respectively, for 0.025 wt.% MWCNT resin compared to unfilled resin. When 0.1 wt.% MWCNT resin was used, tensile stress increased by ~7.5%, while fracture stress increased by ~33% compared to pure resin. While mechanical performance of specimens was improved, the addition of MWCNTs resulted in more brittle material compared to pure resin. The specimens fabricated with MWCNTs showed a decrease in elongation at break of ~28%.

The specimens fabricated with MWCNTs showed a decrease in elongation at break of ~28%. Also, it was noted that increasing the loading of the MWCNTs from 0.025% to 0.1% did not significantly affect mechanical properties and brittleness.Athreya, et al. [58] used nanosized carbon black (CB) powder blended with Nylon-12 powder for Laser Sintering (LS). Their work was focused on finding optimal printing conditions of the nanocomposites. Flexural modulus of Nylon-12/CB composites was found to be lower than pure Nylon-12, depending on scan speed and laser power. The “optimum” parameters were found to be 4W laser power with 30 in/sec scan speeds. Also, it was observed that some portion of particles do not melt during the SLS process, but they remain in their original form in the final printed part. Electrical conductivity of final parts was five orders of magnitude higher for parts printed from Nylon-12/CB composite compared to pure Nylon-12.Athreya, et al.

Electrical conductivity of final parts was five orders of magnitude higher for parts printed from Nylon-12/CB composite compared to pure Nylon-12.Athreya, et al. [59] also compared Nylon-12/CB parts made by LS with objects made by extrusion-injection molding. As in [58], Nylon-12/CB composites were compared with pure Nylon-12. It was found that parts prepared by LS have lower tensile modulus than those prepared using molding. Pure Nylon-12 parts prepared by molding had higher impact strength compared to parts fabricated by LS. The addition of CB to the Nylon-12 resulted in a decrease in impact strength for both LS and molding parts. Notable results were observed in the dispersion of CB within the polymer after printing. It was shown that LS led to formation of large agglomerates of CB within the polymer, whereas parts prepared by the extrusion molding demonstrated uniformly distributed sub-100nm aggregates.

It was shown that LS led to formation of large agglomerates of CB within the polymer, whereas parts prepared by the extrusion molding demonstrated uniformly distributed sub-100nm aggregates. Furthermore, the porosity of parts produced by LS was ~10%, while parts fabricated by extrusion had porosity of <5%.Goodridge, et al., [60] used carbon nanofibers (CNF) as additives to Polyamide-12 for LS. Printing material was prepared by melt mixing 3 wt.% CNF with Polyamide-12. Three systems were compared: 1) as-supplied Polyamide-12 powder, 2) CNF/polyamide-12 (melt mixing), and 3) Polyamide-12 with all the steps of preparation of the CNF/polyamide-12 mixture, but without CNF (Neat Polyamide-12). Figure 5 shows the photographs of the powderbed and printed parts for the three materials. Sintered parts prepared using CNF/Polyamide-12 and pure Polyamide-12/melt mixing had a rough surface compared to as-received Polyamide-12.

Sintered parts prepared using CNF/Polyamide-12 and pure Polyamide-12/melt mixing had a rough surface compared to as-received Polyamide-12. Storage modulus (demonstration of elastic response) and loss modulus (viscous response) were determined for all three systems. Parts printed with as-supplied Polyamide-12 demonstrate the highest moduli compared to the other two systems. Parts prepared using CNF showed 22% and 15% higher storage modulus and loss modulus, respectively, compared to Polyamide-12/melt mixing. SEM analysis showed that CNFs are well dispersed in the polymer after the LS procedure. Overall, the addition of CNFs improved the mechanical properties of sintered parts compared to similarly prepared Polyamide-12, but the resulting parts were not smooth.Figure 5. Photographs of the powder bed (top) and fabricated parts (bottom) using three different materials as indicated. Ref. [60]Similar studies were performed by Lao, et al. [61].

Photographs of the powder bed (top) and fabricated parts (bottom) using three different materials as indicated. Ref. [60]Similar studies were performed by Lao, et al. [61]. They used MWCNT melt-blended with Polyamide-11. Injection and compression molded specimens were prepared for characterization of their mechanical and thermal properties. Tensile strength and tensile modulus increased with increasing MWCNT loading, while elongation at break decreased with increasing MWCNT loadings. Thermo-gravimetric analysis (TGA) analysis showed that thermal stability increases with increase in MWCNT loading. At 50% mass loss, decomposition for 7 wt.% temperature MWCNT/Polyamide mixture.

Thermo-gravimetric analysis (TGA) analysis showed that thermal stability increases with increase in MWCNT loading. At 50% mass loss, decomposition for 7 wt.% temperature MWCNT/Polyamide mixture. The electrical resistivity of polymer nanocomposites decreased with increase of MWCNT loading.increasedfrom 439℃for pure Polyamide-11to 488℃Yildirim, et al., [62] demonstrated the excellent biocompatibility of Single Walled Carbon Nanotubes (SWCNT) – alginate composites for application as a tissue scaffolds. Tissue scaffolds were prepared by Freeform Fabrication technique (an analog of FFF), using 1 wt.% SWCNT mixed with alginate as a printing material. Figure 6 shows the morphology of printed three-dimensional scaffolds with and without SWCNT. SWCNTs were evenly dispersed on the alginate surface, but the printed surface was rough compared to pure alginate.

SWCNTs were evenly dispersed on the alginate surface, but the printed surface was rough compared to pure alginate. Mechanical testing showed higher tensile stress and tensile modulus for alginate/SWCNT composites, but smaller elongation at break values compared to pure alginate. More importantly, it was demonstrated that alginate/SWCNT composites are promising materials for tissue growth. Scaffolds printed using SWCNTs showed a constant cell proliferation rate for up to one weekand the number of cells on day seven was six times higher than on day one; the rate of proliferation is higher than pure alginate scaffolds.Figure 6. Photographs (left side) and micrographs (right side) of fabricated 3D scaffolds using (A) SWCNTs/Alginate composite and (B) pure alginate. Ref. [62]These experiments demonstrate that the addition of carbon nanotubes to SL resins and LS materials can improve mechanical properties of final parts.

Ref. [62]These experiments demonstrate that the addition of carbon nanotubes to SL resins and LS materials can improve mechanical properties of final parts. In all cases, tensile stress and fracture stress increased with the addition of carbon nanomaterials, while elongation at break decreased. The addition of carbon nanomaterials also increased electrical conductivity of the final parts. Finally, it was shown that carbon nanotube–alginate tissue scaffolds can improve the rate of cell proliferation during culture tests.Ceramic and Semiconductor NanoparticlesCeramics are usually inorganic, non-metallic materials such as oxides, carbides and nitrides. As in the case of metal NPs, the size of ceramic NPs plays a crucial role, especially on sintering temperature. Some ceramic materials such as SiO2 and TiO2 show semiconducting properties. A semiconductor is a material with electrical conductivity in-between a conductor and an insulator.

Some ceramic materials such as SiO2 and TiO2 show semiconducting properties. A semiconductor is a material with electrical conductivity in-between a conductor and an insulator. Such materials are characterized by conduction and valence energy bands, which are separated by a known energy range (band gap). The band gap affects the semiconductor properties, such as optical and electrical properties. By altering the size of semiconductor nanoparticles, one can change the band gap and modify the properties of the material. Extensively studied semiconductor nanoparticles include cadmium chalcogenides (Quantum Dots) [63, 64]. However, other types of semiconductor nanoparticles have been synthesized and investigated [65-69]. Figure 7 shows the photographs of colloidal solutions of Cd free Quantum Dots. These materials have found application in optoelectronic devices [70],solar cells [71], bio-imaging [72, 73], and labeling [74].

These materials have found application in optoelectronic devices [70],solar cells [71], bio-imaging [72, 73], and labeling [74]. Here we describe the effects of the addition of ceramic and semiconductor nanoparticles on mechanical properties of parts fabricated by AM.Figure 7. Brightly glowing vials of highly luminescent, water soluble quantum dots produced by the new NIST microwave process span a wavelength range from 500 to 600 nm. Credit: NIST (National Institute of Standards and Technology).Chung, et al. [75] used Nylon-11 loaded with silica NPs in LS to create Functionally Graded Material (FGM) composite components. FGM materials combine different composition and properties within a single component to optimize performance of that component. Silica loadings in Nylon-11 investigated were 0, 2, 4, 6 and 10 wt.%. The particle size of Nylon-11 powder was 106 to 150µm, while the size of silica NPs was ~15nm.

Silica loadings in Nylon-11 investigated were 0, 2, 4, 6 and 10 wt.%. The particle size of Nylon-11 powder was 106 to 150µm, while the size of silica NPs was ~15nm. Adjustment of printing parameters led to part density of >90% for different silica loadings. Transmission electron microscope (TEM) analysis confirmed a homogeneous distribution of silica particles inside the polymer. Mechanical tests showed that with increase of silica loading, the parts become stiffer (with an increase in tensile modulus), but also more brittle (with a decrease in tensile strain at break). It is expected that there is a critical composition at which these characteristics reverse these trends.Cai, et al. [76] demonstrated Direct Writing construction of three-dimensional meshes from TiO2 – vaseline composites.

[76] demonstrated Direct Writing construction of three-dimensional meshes from TiO2 – vaseline composites. Vaseline was loaded with 0, 2.5, 5, 7.5, and 10 wt.% of TiO2 NPs (diameter = 21nm); ink properties were investigated; 3D scaffolds were fabricated; and mechanical performance of printed scaffolds was tested. Creep and recovery tests showed that inks with low NP content (0-5 wt.%) form structures that easily damage and don’t recover, while inks with high NP loading demonstrated a tendency toward lower deformation and increased elastic properties. Fabricated inks showed an excellent ability to construct 3D structures. An increase in NP content reduced rod bending and caused shrinking and folding. Scaffolds produced on the flexible substrate exhibited an ability to preserve their shape after repeated bending and stretching.Figure 15.

Scaffolds produced on the flexible substrate exhibited an ability to preserve their shape after repeated bending and stretching.Figure 15. Images of assembled structures on different scale: (a) a four-layer scaffold piled by perpendicular rods with diameter of 500 μm and inter-rod distance of 2 mm in a 20 mm × 20 mm square, (b) bending of the formed structure, (c) agglomeration of nanoparticles at the rod’s surface, (d) magnified image of nanoparticles doped ink. Ref. [76]Semiconductor nanoparticles were used in SL to influence the final mechanical characteristics of the fabricated parts by Duan et. al. [77]. The addition of TiO2 NPs (diameter=30nm) to the photosensitive resin showed promising results for improving the mechanical properties of fabricated parts. NP loadings from 0.1 to 5.0 wt.% were used. Viscosity of fabricated resins increased with increasing NP content. Properties of the final parts were studied extensively on 0.25 wt.% NP loading.

NP loadings from 0.1 to 5.0 wt.% were used. Viscosity of fabricated resins increased with increasing NP content. Properties of the final parts were studied extensively on 0.25 wt.% NP loading. The addition of TiO2 NPs improved the mechanical properties of the final part; tensile strength increased by 89%, tensile modulus increased by 18%, flexural strength increased by 6%, and hardness increased by 5%. All these parameters indicate the improvement of plasticity of fabricated parts. Additionally, application of NPs improved thermal stability of the fabricated parts.The Lewis group [78] used BaTiO3 NPs (diameter=60nm) coated with poly (acrylic acid) to fabricate inks with high (>45%) solid loadings for direct writing. Experiments showed that the addition of NPs leads to an increase in shear stress and elastic modulus by one order of magnitude as the result of fluid-to-gel transition. This transition was performed using mono- or divalent salt species.

This transition was performed using mono- or divalent salt species. It was demonstrated that the nature of salt used for fluid-to-gel transition greatly affects the elastic modulus. Elastic modulus was approximately two orders of magnitude higher in the case when divalent salt was used. Successful demonstration of printing of 2D and 3D structures with feature sizes of ~30µm and ~100µm was performed.Experimental demonstration of the adjustment of the properties of printed parts by surface modification of NPs was offered by Zheng, et al. [79]. Al2O3 nanoparticles (diameter=60nm) were used in combination with polystyrene (PS) for LS. Three compositions were studied: pure PS, Al2O3/PS mixture, and Al2O3/PS core-shell particles produced by emulsion polymerization. Sintering of Al2O3/PS core-shell particles led to formation of acompact structure, while Al2O3/PS mixture developed highly porous surface with numerous voids.

Sintering of Al2O3/PS core-shell particles led to formation of acompact structure, while Al2O3/PS mixture developed highly porous surface with numerous voids. Encapsulating NPs into the polymer matrix led to uniform dispersion of the NPs in the matrix, while NPs mixed with PS tended to form aggregates. Mechanical properties of all three materials were investigated. It was shown that impact strength and tensile strength increased (300% and 50%, respectively) for Al2O3/PS core-shell specimens, compared to Al2O3/PS mixture and pure PS. Al2O3/PS core-shell samples were more ductile and demonstrated higher toughness.The potential use of nanocomposite scaffolds for bone tissue engineering was demonstrated by Duan, et al. [80]. Calcium phosphate (Ca-P) and carbonated hydroxyapatite (CHAp) NPs were incorporated into poly(hydroxybutyrate-co-hydroxyvalerate) (PHBV) and poly(L-lactic acid) (PLLA), respectively. The size of calcium NPs varied from 10 to 30 nm.

The size of calcium NPs varied from 10 to 30 nm. These nanocomposites were used to fabricate tissue scaffolds using the LS technique. Mechanical properties, such as compressive strength and modulus, of nanocomposite scaffolds were higher than those of pure polymers in dry environment. Biological evaluation of fabricated scaffolds was carried out with human osteoblast-like cells (SaOS-2). After three days of culture growth, almost no dead cells were found on all four types of scaffolds, thus indicating high cell viability of the fabricated scaffolds. Cell proliferation was enhanced in the case of Ca-P/PHBV scaffolds compared to pure PHBV, while no difference was observed in cell proliferation between CHAp/PLLA and pure PLLA scaffolds.

Cell proliferation was enhanced in the case of Ca-P/PHBV scaffolds compared to pure PHBV, while no difference was observed in cell proliferation between CHAp/PLLA and pure PLLA scaffolds. After seven days, cultured cells were well attached and spread over the scaffold surface for all four types of materials.In summary, the use of semiconductor and ceramic nanoparticles as additives to LS, SL, and Direct Writing materials has shown promising results in enhancing mechanical properties of final printed parts. Fabrication of FGM components and tissue scaffolds has been also successfully demonstrated. Biodegradable ceramic nanostructures can be successfully applied in bone tissue engineering.Summary of Published LiteratureTable 1 summarizes the state of the art on the inclusion of nanomaterials in AM as of thiswriting.Table 3.

Summary of published literature of AM with nanomaterials.743Challenges in AM and NanomaterialsThere are several challenges inherent in the merger of AM and nanomaterials. The most important issue that needs to be resolved is agglomeration of the nanomaterials in printing media when working with extrusion methods, 3DP, and Stereolithography. Agglomeration of the particles most likely will lead to loss of their unique properties (i.e., nanomaterial will behave as a bulk). To overcome this issue, functionalization of the nanomaterials with organic linker molecules is required. Linkers can keep particles away from each other, even when embedded into the printing media.

To overcome this issue, functionalization of the nanomaterials with organic linker molecules is required. Linkers can keep particles away from each other, even when embedded into the printing media. Solving this problem will lead to the design of wholly new nanocomposite inks that can improve the properties of the final parts and preserve the unique characteristics of the nanomaterial embedded into the printing material.When using SL in combination with nanomaterials, one should find suitable nanomaterials for a given wavelength of the UV light source. The cure depth of the SL can be affected by the presence of nanomaterials due to their ability to absorb the UV light. Much research will be required to overcome this challenge.

The cure depth of the SL can be affected by the presence of nanomaterials due to their ability to absorb the UV light. Much research will be required to overcome this challenge. Experiments on finding the ideal wavelength that will solidify the polymer without being affected by the presence of the nanostructures may be necessary.Published work on inclusion of nanomaterials into LS, in some cases, showed that porosity of the final printed parts is higher compared to parts printed without nanomaterials. This could be due to the different nature of materials used (carbon black and Nylon-12, or Silica NPs and Nylon-11, etc.). The synthesis of core-shell nanostructures, where the core consists of the nanomaterial of interest and the shell consists of the printing material, can improve the density of the final parts [79].

The synthesis of core-shell nanostructures, where the core consists of the nanomaterial of interest and the shell consists of the printing material, can improve the density of the final parts [79]. Resolving this issue will provide new printing materials that possess unique properties – electrical and thermal conductivities, hardness – while simultaneously yielding nanocomposites as dense as nanostructure-free printed parts.3DP and extrusion processes also have their challenges, when combined with nanomaterials. The most significant issue is nozzle clogging when high loadings of nanomaterials are used. This could be due to the formation of aggregates within the printing material. More experimental work is required to determine the ideal composition of nano-inks, that could flow freely through nozzles.ConclusionsThis review provides an overview of the work done to date to incorporate nanostructures into Additive Manufacturing (AM).

The addition of metal nanoparticles generally decreases sintering temperatures, improves part density, and decreases shrinkage and distortion of printed parts. Metal nanoparticles embedded into polymer materials can also provide improved electrical conductivity in fabricated objects. Incorporation of carbon nanotubes in printing media offers a potential route to improving mechanical properties of the final parts and to increasing electrical and thermal conductivities. The addition of carbon nanotubes in bio-scaffolds can yield excellent enhancement of cell proliferation. Adding semiconductor and ceramic nanoparticles to printing media can lead to improvements in mechanical properties of the final parts. Ceramic nanoparticles can be effectively used for bone tissue engineering.Despite these early successes, challenges in the application of nanomaterials to AM are nevertheless numerous.

Ceramic nanoparticles can be effectively used for bone tissue engineering.Despite these early successes, challenges in the application of nanomaterials to AM are nevertheless numerous. Each of the AM methods described has its own inherent limitations (nozzle clogging, aggregation within printing media, rough surface finish of printed parts etc.)when nanoparticles are applied with the respective printing media. Overcoming these design boundaries may require the development of new instrumentation for successful AM with nanomaterials.It is shown that there are many opportunities in the marriage of AM and nanotechnology. By leveraging the tunable properties of nanostructures, we can expand the material properties, and thus applications, of printed components. Moreover, it may be possible with AM-produced nanocomposites to create objects with graded material properties by varying the nanostructure loadings during a part’s synthesis.

Moreover, it may be possible with AM-produced nanocomposites to create objects with graded material properties by varying the nanostructure loadings during a part’s synthesis. Combining multiple nanomaterials in the same AM part would then allow us to move from simple to more complex printed objects such as fuel cells, batteries, solar cells, etc. Recent advancements in nano-biomaterials could also further enable the printing of replacement organs and bone. Again, these forthcoming directions will require not only a greater focus on the development of new materials for existing AM systems, but they also may necessitate the development of wholly new AM approaches.Overall, we find promising results have been published in the application of nanomaterials and AM, yet significant work remains to fully harness their inherent potential.

ditive Manufacturing Education in the UKPhill Dickens – Loughborough University Phill Reeves – Econolysts Ltd. Richard Hague –University of NottinghamIntroductionThere has been and continues to be a considerable effort regarding education for Additive Manufacturing (AM) in the UK. This generally started in 1992 with a seminar organised by the Institution of Mechanical Engineers and an industrial exhibition stand at Mach 92.

This generally started in 1992 with a seminar organised by the Institution of Mechanical Engineers and an industrial exhibition stand at Mach 92. However, before the education activities are discussed it is useful to show the AM research landscape in the UK as this will give an indication of the level of activity It should be noted that there will also be a considerable number of organisations involved in using AM but not involved in research.The UK landscape for AM ResearchThis paper contains statistical data derived from a study undertaken by Econolyst Ltd for the Additive Manufacturing (AM) Special Interest Group (AM-SIG), coordinated by the Materials Knowledge Transfer Network (KTN) at the request of the UK’s Technology Strategy Board (TSB).

As such it should be considered as a ‘snapshot’ of the UK AM research community in Q1 of 2012 and looks at the period from 2006 to 2012.Overview of UK AM research community81 organizations within the UK have been identified as currently or previously (back to 2007) engaging in AM research activities.

This includes 24 Universities and 57 companies, as detailed in Figure 1.Figure 1 - List of organizations included in AM SIG research review processUniversities & Research Technology Organizations (RTO’s)Bath University Birmingham City (JIIC) Birmingham University Cambridge University Cardiff University CERAM Cranfield DMUExeter Glasgow Caledonian Lancaster Lboro Leicester Liverpool London metropolitan ManchesterMaterials Solutions Newcastle Nottingham Sheffield University TWI University of East Anglia Warwick Wolverhampton1Within these 81 organizations, 151 AM machine platforms have been identified, including 109 commercial machine tools procured from technology vendors, and 46 machines either modified from commercial machines (for specific research tasks), or self-assembled using modular elements such as robots, gantries, lasers and material feed nozzles.

A detailed breakdown of machines by technology class within the UK can be seen in Figure 2.Figure 2 - Distribution of AM machine platforms in the UK used for researchOf the installed machine base within the UK, 52% of University machine capacity is used directly for research.

This compares to just 38% within industry, where machines are used to support prototyping and low volume production applications.2A detailed analysis of machine usage within the organizations canvased in this research can be found in Figure 3.Figure 3 - AM technology usage profile for industrial and research organizations3The largest percentage of employment in both academic and industrial establishments is at a post graduate or post-doctoral level, as opposed to a technician level, as detailed in Figure 4.Figure 4 - Typical distribution of staff within UK AM research facilitiesAM education is cascading down to an undergraduate level, with 17 universities across the UK engaged in the delivery of AM related teaching or training courses, ranging from dedicated post-graduate courses at De Montfort and Wolverhampton Universities, though to under graduate modules delivered at 10 Universities up and down the country.

It should be noted however that as yet, there are no dedicated Additive Manufacturing or 3D Printing undergraduate offerings. A detailed breakdown of AM teaching and training provision within the UK can be found in Figure 5.4Figure 5 - AM teaching & training activity within the UKInvestment into AM Research and Technology Transfer within the UKBetween 2007 and 2016, £95.6-million has either been invested, or is committed to be invested in UK Additive Manufacturing research & technology transfer activities within the organizations listed in Figure 1. This includes £80-million of research focused monies and £15.6-million of funding for technology transfer activities.

This includes £80-million of research focused monies and £15.6-million of funding for technology transfer activities. Of this funding, the largest proportion, £24.9-million has come directly from industry, with approximately £13-million contributed each by the TSB, EU Framework Programs (FP6 & FP7), and the European Regional Development Fund (ERDF).5Investment in industrial AM R&D activitiesBy sector, aerospace has received the largest grant investment to support AM research (£20.5-million to industry & universities combined), albeit the sector has also made the largest private sector direct investment (£13-million), as detailed in Figure 6. The medical sector has been the second largest beneficiary of research funding (£12- million), but has contributed only £3-million directly to AM research activities through matched funding. This is slightly less than the automotive sector who have invested some £3.5-million, but which has only received grant support of £7-million.

This is slightly less than the automotive sector who have invested some £3.5-million, but which has only received grant support of £7-million. In short, aerospace has invested in and benefited from AM research beyond any other sector.Figure 6 - AM R&D investment by sector & sources (2007 - 2016)Further analysis shows the split of R&D investment between industry and academia, based on grant and match funding together. Aerospace businesses have received a similar level of direct grant funding to the investment they have made. The creative industries and the fashion sectors on the other hand have received very little funding directly, with the majority of grant funding going directly to research organizations such as Universities.

Although only accounting for a small proportion of research activity, the oil, gas & chemicals sectors and the power generation sector have all received direct funding relative to their industry investment, at a rate on par with the aerospace sector.6In summary, they have benefited from similar levels of funding, but have engaged in a much lower level of research activity.In terms of direct investment by the private sectors across the whole AM supply chain, aerospace dominated the applications / user domain, driven by companies including Boeing, Bombardier, EADS, Messier Dowty, Rolls Royce and Aero Engine Controls.

After the aerospace sector, as detailed in Figure 7, the greatest investment within the UK has come from the software sector, largely driven by companies including Delcam, Granta, Materialise UK and Simpleware.Figure 7 - Investment in AM R&D by supply chain positionUniversity and RTO investmentIn the UK AM research community (Universities and RTO’s), £51.4M has been invested in R&D activities and £16.2M in technology transfer (2007–2016 time frame).The UK within the EUThere are currently 20 active EU FP7 projects with work packages focused on Additive Manufacturing. A further 2 projects have finished in recent months, with one FP6 project dedicated to AM finishing in 2010. The current active FP7 projects are worth a7combined €99.3-million of public and private sector investment, of which €84-million relates directly to AM aligned work packages.Of the 20 current projects, the UK leads nine projects (45%), this being more than any other country in Europe.

Moreover, of the 240 partners engaged in collaborative AM projects around Europe at this time, 54 are UK based organizations, again representing the largest proportion of participants by any EU member state (23%).One alternative to measuring financial input to research activity is to measure one of the outputs, namely, research publications. This allows us to identify a number of important metrics, including the locations of research critical mass, and the focus of research activity in terms of technologies and applications.To undertake this review it was decided to focus on conference base publications. This decision was taken given the limited number of dedicated AM related journals and peer reviewed publication titles. A review was therefore undertaken of all AM related conferences taking place in 2011, including international conferences known to have sessions dedicated to AM activities. A list of the 17 conferences identified can be seen in Figure 8.

From these conferences a review was undertaken of 495 individual papers or abstracts.Figure 8 – List of AM related conferencesConferenceLocation (Country)Number of PapersConference focus3DSUG AFPR AMIC (Loughborough) ICAT NZRPD RAPDASA RapidTech RPDM RPS SFF SME Rapid TCT VRAP WohlersUSA France UK Slovenia New Zealand South Africa Germany UK Japan USA USA UK Portugal Germany24 32 20 42 14 37 37 17 18 56 51 21 85 10AM & RP AM & RP AM AM AM & RP AM & RP AM AM & RP AM & RP AM AM & RP AM & RP AM & VR AMDigiMan Powders AUS TENGUSA Australia Japan10-relevant 8 - relevant 13 - relevant 495 TotalDigital printing Powders Tissue engineering8As shown in Figure 9, 54% of all Additive Layer Manufacturing conference papers were authored in the EU, with 26% authored in the USA and the remaining 20% authored in the rest of the world.Figure 9 - Percentage distribution of AM conference papers by regionIf we then consider these statistics at an individual country level, as shown in Figure 10, we can see that the USA accounts for the largest number of papers (129), followed by the UK (75) and then Germany (71).Figure 10 - Number of AM conference papers by leading countriesOf course, if we consider these statistics in terms of per capita population, we see a very different picture.

Using 2010 World Bank data on population (USA = 311,591,917 / UK = 62,218,761 / Germany 81,702,329), we find that the UK produces 315% more AM related paper per capita than the USA, and 40% more than Germany. This would clearly suggest that the UK AM research community holds a prominent global position. It could of course be argued that UK conference focus on UK papers and presentation and therefore slew these results.9However, this is at odds with online usage of the Rapid Prototyping Journal. Between 2005 and 2011, (online) usage in the UK increased by 273% whereas for the same period, usage worldwide increased by 419% (and in the USA specifically this increase was higher at 426%).There has been a very strong growth in the numbers attending the TCT exhibition as shown in Figure 11.

In the early years the emphasis was on the conference whereas later years concentrated on the exhibition.Figure 11 TCT Exhibition AttendanceTCT Exhibition (UK)60005000400030002000100001997 1998 1999 2000 2001 2002 2003 2004 2005 2006 2007 2008 2009 2010 2011However, the attendance at Additive Manufacturing International Conference (AMIC) (Loughborough) has also increased as shown in Figure 12.10Figure 12 Attendance at AMICAttendance at AMIC (Loughborough)2502001501005002007 2008 2009 2010 2011 2012In summary, it can be concluded that the UK is unarguably one of the world’s leading sources of AM related knowledge and research activity, along with Germany and the USA.However, this now raises a number of questions such as:What education activities have taken place?  How many people have been involved in receiving education and to what level?  What benefit have these people obtained from the education?  What benefit has there been to UK plc?  Which education activities have provided the most benefit?  Which education activities should we drop and which ones should we increase?  Was all this knowledge creation worthwhile?11AM Education activitiesThe education regarding Additive Manufacturing in the UK really started in earnest during 1992.

During this year there were a number of seminars and conferences including the 1st European Conference on Rapid Prototyping which was held at University of Nottingham.During the past 20 years the education activities have included a variety of print media, television programmes, web sites, events, research projects, technology transfer, formal teaching etc. This work has been predominantly undertaken by universities, service bureaus, machine and material suppliers and media organisations.

It would be very difficult to calculate the number of person years involved but it is likely that it would be a large number.Examples of activities can be seen in Figure 13.PrintExamplesResearch JournalsMagazinesRapid Prototyping Journal Prototyping Technology International Precision ToolmakerEngineering DesignerCastings BuyerPro/E MagazineTCTCadDesk MagazineEconomistNewspaper articles BooksTelevisionOnlineEventsWeb sitesConferencesExhibitionsSeminarsTomorrows World Machine sellers University European Conference on Rapid Prototyping 1992 -MachIMechE Seminar on Rapid Prototyping Systems, 1992Life on EarthBureausYoutube National Conference on Rapid Prototyping & Tooling Research, 1995 to date Celebration of UK Manufacturing (EPSRC) DTI Seminar 1992Discovery Channel ConsultantsLoughborough Conference 2006 to dateScience MuseumRapid Prototyping & Manufacturing Association Seminar, 2001 -Sky NewsCasting companiesTCT Conference 1996 to dateTCT LiveAssociation of Laser Users12Ongoing ActivitiesResearch ProjectsEUEPSRCDTI/TSBDept of HealthTechnology Transfer ActivitiesInforming Stakeholders University Courses1993 - Technology Transfer Centre for Rapid Prototyping, University of Nottingham TSB/BER/BIS Visits Masters Training Programme in Rapid Product Development.Rapid Manufacturing ConsortiumH&S VisitsUndergraduate courses and away)(homeForesight TeamConsulting Activities Community BuildingNational Centre for AMUnfortunately there is very little information available on the number of people involved in the activities above.

Even if information was kept at one time it has largely been lost. However, it is probably more important to quantify the benefit to individuals and here there is no real information. Therefore, it is not possible to determine what benefit has accrued to the UK economy, which activities have been most useful and would be in the future.It is clear that there is real potential here for a research project to try and track this activity properly and the resulting benefits.AcknowledgementsMuch of the information in this paper was supplied by Econolyst Ltd., TCT and Rapid Prototyping Journal.13.

ternational SFF Symposium University of Texas at Austin August 12-14, 2013, Austin, TexasAdditive Manufacturing Laser Deposition of Ti-6Al-4V for Aerospace Repair ApplicationsN.K.Dey*, F.W.Liou*, C. Nedic**Department of Mechanical & Aerospace Engineering Missouri University of Science & Technology, Rolla**GKN Aerospace, Hazelwood, St Louis, MOABSTRACTParts or products from high performance metal are very expensive, partly due to the processing complexities during manufacturing. The purpose of this project is to use additive laser deposition and machining processes to repair titanium parts, thus extending the service life of these parts. The study broadly laser deposition, tests. machining, Comparative study of mechanical properties (UTS, YS, percentage elongation) of the repaired samples to the ideal conditions was undertaken.

The study broadly laser deposition, tests. machining, Comparative study of mechanical properties (UTS, YS, percentage elongation) of the repaired samples to the ideal conditions was undertaken. The research throws up interesting facts where the data from the test sample shows enhancement of properties of the repaired part.included preparingthe defects,sample preparationand mechanical1. INTRODUCTION In recent years, the significance of using laser metal powder deposition for repair has increased manifolds. This research paper focuses on development of a reliable repair process of damaged Ti-6Al-4V parts using hybrid manufacturing technology at in house Laser Aided Manufacturing Process (LAMP) Laboratory. The basic principle of the laser process is a laser beam creating a metallurgical bond between the damaged surface area of a part, known as the substrate, and the added repair material, which is a metal powder [1].

The focused laser beam melts the surface of the target material and generates a small molten pool of base material. Powder delivered into this same spot is absorbed into the melt pool, thus generating a deposit that may range from 0.005 to 0.040 in. thick and 0.040 to 0.160 in. wide. The resulting deposits may then be used to build or repair metal parts for a variety of different applications. The economic advantage of this process is the value addition to the component by layer by layer deposition of material that significantly increases the operating life of the component that needs to be run reliably over an extended period of time. 8531. INTRODUCTION In recent years, the significance of using laser metal powder deposition for repair has increased manifolds. This research paper focuses on development of a reliable repair process of damaged Ti-6Al-4V parts using hybrid manufacturing technology at in house Laser Aided Manufacturing Process (LAMP) Laboratory.

The basic principle of the laser process is a laser beam creating a metallurgical bond between the damaged surface area of a part, known as the substrate, and the added repair material, which is a metal powder [1]. The focused laser beam melts the surface of the target material and generates a small molten pool of base material. Powder delivered into this same spot is absorbed into the melt pool, thus generating a deposit that may range from 0.005 to 0.040 in. thick and 0.040 to 0.160 in. wide. The resulting deposits may then be used to build or repair metal parts for a variety of different applications. The economic advantage of this process is the value addition to the component by layer by layer deposition of material that significantly increases the operating life of the component that needs to be run reliably over an extended period of time. 853Aerospace parts are generally very complex and expensive.

853Aerospace parts are generally very complex and expensive. It is found to be more economical to repair the parts rather than procuring new parts, as the new parts are either exorbitantly expensive or require a large lead time for delivery. In some cases the original equipment manufacturer might be out of business. Under is a more viable and economical option for getting the parts restored to their original form.these circumstances repairConventional welding processes (i.e. tungsten inert gas or gas metal arc) create heat affected zones in the substrates close to the weld. This often causes distortion or change in properties, which makes the repair process even more difficult. Moreover, the metallurgical properties of the heat affected zone can be different from the base material far away from the weld. Laser deposition has a small and limited heat affected zone which makes the weld less vulnerable to degradation of properties.

Laser deposition has a small and limited heat affected zone which makes the weld less vulnerable to degradation of properties. Further, using this process a better microstructure and a smaller dilution zone can be achieved. This allows repeatability and reproducibility of layer by layer deposition of material. The powder filler material is fed by a nozzle in an inert gas atmosphere, as powder gives us more flexibility of chemical composition of filler material.for greater2. CONCEPT AND PROCESS The research objective is to use laser deposition and machining processes to repair damaged titanium parts of alloy Ti-6Al-4V (Ti64). The ability to repair Ti64 parts with sound microstructure using metal additive processes is currently at nascent stage. This effort is a significant enabler to the metal repair technology by using hybrid laser deposition and machining technology that is currently rudimentary and limiting implementation of the technology.

Ti64 is essentially alpha-beta structured Titanium alloy which have an excellent combination of strength and ductility. They are stronger than the alpha or beta alloys. Welding alpha-beta alloy titanium can significantly change strength, ductility and toughness characteristics as a result of the thermal cycles to which they are exposed. The change in the ductility can be attributed to the phase transformation of the alloy. Ti64 has an excellent weldability, which can be attributed to two principle factors. First, the alpha-prime martensite that forms in the Ti64 is not as hard and brittle as that exhibited by more heavily beta- stabilized alloys. Secondly, Ti64 exhibits a relatively low hardenability on cooling from solution treatment temperatures. This behavior allows the formation of higher proportion of the beta more microstructure even at relatively high rates of weld cooling.

This behavior allows the formation of higher proportion of the beta more microstructure even at relatively high rates of weld cooling. Ti64 when welded with filler metal of matching composition of extra low interstitial grade, it improves the ductility and toughness. Due to its single phase mode of solidification (i.e. absence of low melting point eutectics), Ti64 is also highly resistant to solidification related and liquation related cracking [2].desirable Widmansttenalpha-plus-retainedFrom the current historical data available from the aerospace in industry, components. These defects may arise from numerous reason like fatigue to manufacturing. These were further classified into different categories. They broadly are deformation defect, undercut defect, cutter pull out defect and cutter mark defect. The classification were carried out in order to give us a better understanding of the damages and to come up with a reliable repair process.

The classification were carried out in order to give us a better understanding of the damages and to come up with a reliable repair process. The classification follows a common pattern evolved during the gathering of data from abundant damaged samples provided by the industry participant. All such major defects were replicated in test parts provided by the industry participant.there are differenttypes of defects foundFig. 2. Test partA procedure involving metal additive and subtractive processes will be investigated. Experiments to validate the procedure forTi64 samples will be carried out. The work will consist of four tasks:Task 1: Sample preparation process study: a procedure for preparation of repair samples is necessary, such as to clean the damaged surface and to determine or make access for tools to reach the damaged area. This could also involve minimum machining. Step 1.

This could also involve minimum machining. Step 1. Finding optimum places for creating the damages – care should be taken that the laser nozzle path should be clear of any obstruction during laser metal deposition. Step 2. Creating the damage (four types of damages) at the designated places – there was a need to cut down the block into four parts for better fixturing in the Laser deposition and CNC machine. Step 3. Checking for foreign particles left in the damaged area. If necessary, a blasting process was applied for removal of all unwanted particles. Step 4. Tool Path Planning – Tool path was calculated for machining the damaged surface, so that Laser Metal Deposition could be performed for the repair.Fig. 3. Laser deposited partsTask 2: Sample repair via additive technology: to generate the deposition path based on measured data will be investigated. The optimal deposition parameters and setup for Ti64 deposition will also be studied. Step 1.

The optimal deposition parameters and setup for Ti64 deposition will also be studied. Step 1. Path planning for Laser Metal Deposition (LMD). This step required precise planning because the laser nozzle has a limited work envelope. Calculations were needed for the time of deposition and preheat pass.Fig. 4. Test specimen zoneFig. 5. Test specimenTask 3: Sample machining/finishing: Samples were machined to original shape. Step 1. Creating Samples – Next was to cut out samples from the repaired portion of the part. Specimens were chosen in such a way that the test area is the most vulnerable spot of the test part. The aim was to cut out the maximum number of test specimens as possible, in order to reduce the variations in the results. Step 2. CNC machining – Cut out the repaired part from the major block. Optimized the cut-out part to get the maximum number of test samples. Milled the surface from both sides to get the samples. Step 3.

CNC machining – Cut out the repaired part from the major block. Optimized the cut-out part to get the maximum number of test samples. Milled the surface from both sides to get the samples. Step 3. The samples were created at the zone between the bonding of the laser deposition to the substrate, and can be seen in fig 4.Task 4. Sample microstructure: Repair samples will be tested for microstructure analysis, UTS, YTS and % Elongation. Step 1. Tensile Test – From the test specimens, tests were conducted for the UTS, YS, and % Elongation. The values were then calculated for the UTS, YS, and % Elongation.3. EXPERIMENT Laser metal deposition was carried out using a diode laser, which produces a single wavelength laser at power level upto 1000 watts. The system is manufactured at 808 nm and is ideally suited for welding, cladding and surface modification. Its armored 1mm fiber is mounted on a Precitec YC50 laser cladding head. A NI real time controller is used to control the laser.

Its armored 1mm fiber is mounted on a Precitec YC50 laser cladding head. A NI real time controller is used to control the laser. The powder feeder can deliver powder at a rate as low as 0.1 g/min. The whole system is mounted on a Fadal 5 axis CNC machine which is used as the motion driver. The deposition nozzle is mounted on the side of the spindle, which forms a hybrid manufacturing system on a single workstation as shown in Fig 7. A multi-axis planning system (MAPS) was used to generate a laser deposition tool path by slicing the solid model.Fig. 7. Laser deposition process in progressBefore the actual laser metal deposition a pre heat pass was generated at 1000 watts to bring the substrate to a higher temperature in order for the powder to get a better bonding during the actual deposition. The metal deposition was carried out at 800 watts of laser power at powder feed rate of 8g/min.Fig. 8.

The metal deposition was carried out at 800 watts of laser power at powder feed rate of 8g/min.Fig. 8. Dimensions of test specimen (inches)The machining operation was carried out on a 5 axis CNC machining center. The milling operations were performed using 0.25” and 0.5” face milling tools at a feed rate of 24in/min and a spindle speed of 6000rpm. The z axis depth varied at not more than 0.005” per cut. The profile cutting of the specimen was performed using 0.1mm diameter tool at a feed rate of 2 in/min at a spindle speed of 24,000 rpm. The z axis depth was not more than 2mm. The samples were taken out of a transitional zone between the deposited layer and the substrate as shown in Fig. 4Fig.9. Milling operationThe tensile tests were carried out using a MTS universal testing machine. The load varied through a wide range reaching upto a maximum 1600 lb over a time period of 20min. The resultant graph plotted for the UTS and YS is shown in Fig. 10Fig. 10. UTS, YS graph3.

The load varied through a wide range reaching upto a maximum 1600 lb over a time period of 20min. The resultant graph plotted for the UTS and YS is shown in Fig. 10Fig. 10. UTS, YS graph3. TESTS & RESULTS Ultimate tensile strength (UTS), is the maximum stress that a material can withstand while being stretched or pulled before failing or breaking. The UTS gives us a fair idea of the strength of the material as it is the resistance offered by the material to a force tending to tear it apart. The UTS is usually found by performing a tensile test and recording the stress versus strain; the highest point of the stress strain curve is the UTS.

The UTS is usually found by performing a tensile test and recording the stress versus strain; the highest point of the stress strain curve is the UTS. Stress is the ratio of the applied load to the cross-sectional area of an element in the tension and is expressed in pounds per square inch (psi) Stress(σ) = Load(L) Area(A)Strain is a measure of the deformation of the material.Strain(ε) = change in length(ΔL)Original length (L)Yield Strength (YS) is the indication of maximum stress that can be developed in a material without causing plastic deformation. It is the stress at which a material exhibits a specified permanent deformation and is a practical approximation of elastic limit. Offset yield strength is determined from a stress-strain diagram. It is the stress corresponding to the intersection of the stress- strain curve, and a line parallel to its straight line portion offset by a specified strain.

It is the stress corresponding to the intersection of the stress- strain curve, and a line parallel to its straight line portion offset by a specified strain. Offset for metals are usually specified as 0.2%, i.e., the intersection of the offset line and the 0-stress axis is at 0.2% strain. This can be viewed in Fig.10The percentage elongation is the maximum elongation of the gage length divided by the original gage length.Percent elongation = final gage length – initial gage length initial gage lengthThe above parameters were chosen to make a comparative study between the properties of the original substrate and that of the repaired samples, in order for us to determine the reliability of the repair process [3].The tensile strength tests were carried out using the MTS tensile strength test machine. The result for the test for the undercut defect is shown in Table 1. The normal value of Ultimate Tensile Strength (UTS) for a virgin Ti64 substrate is 950 MPa.

The result for the test for the undercut defect is shown in Table 1. The normal value of Ultimate Tensile Strength (UTS) for a virgin Ti64 substrate is 950 MPa. The value for Yield Strength (YS) of a Ti64 substrate is 880 MPa. The normal percentage elongation (% elongation) is 14%. The substrate data gives us a fair idea of our test result.Table 1. UTS, YS, and % elongation of undercut defectTable 2 shows the result from the cutter pull out defect. We can see that the value for the UTS, YS, and % elongation is on the higher side than that of the virgin substrate. Table 3 shows the result for the deformation defect. Here the results are very close to the data for the virgin substrate.Table 2. UTS, YS, and % elongation of cutter pull out defectTable 3. UTS, YS, and % elongation of deformation defectTable 4 shows a comparative study of the UTS of the various samples, from all types of defects, to the value of the substrate.

UTS, YS, and % elongation of deformation defectTable 4 shows a comparative study of the UTS of the various samples, from all types of defects, to the value of the substrate. We can see that the average values of UTS for the deformed defect and undercut defect are very close to that of the substrate. On the other hand the average value of the cutter pull out defect is slightly higher than that of the substrate.Table 4. Comparative study UTS resultsTable 5 shows a comparative study of the YS of the various samples from all types of the defect to the value of the substrate. We can see that the average values of YS of deformed defect and undercut defect is very close to that of the substrate. On the other hand the average value of the cutter pull out defect is slightly higher than that of the substrate. This is very much similar to the result of the UTS comparative study.Table 5. Comparative study of YS resultTable 6.

This is very much similar to the result of the UTS comparative study.Table 5. Comparative study of YS resultTable 6. Comparative study of % elongation resultTable 6 below shows a comparative study of the % elongation of the various samples from all types of defects, to the value of the substrate. We can see that the average value of % elongation for the cutter pull out defect is very close to that of the substrate. On the other hand the average value of the deformed defect and undercut defect are slightly higher than that of the substrate. We can also see that the variation in the data is slightly more than that of the other comparative study.4.1 DISCUSSION OF RESULT The ideal properties for Ti-6Al-4V substrate are as follows:UTS – 950 MPa YS - 880 MPa % Elongation – 14 %The average value of the UTS and YS from the samples containing an undercut defect is 946 MPa and 881.63 respectively [Table 1], which is very close to the ideal value.

The average value of the % elongation [Table 1] from the samples containing the undercut defect is 19.34%, which is on the higher side. This shows an increase in the ductility of the samples. The average value of the UTS and YS from the samples with the cutter pull out defect is 1005.13 MPa and 981.88 [Table 2] respectively, which is on a higher side of the ideal values. This shows an enhancement of mechanical properties in the samples. The average value of the %elongation [Table 2] from the samples of the cutter pull out defect is 15.87%, which is veryclose to the ideal value. The average value of the UTS and YS from the samples containing a deformation defect is 967 MPa and 898 [Table 3] respectively, which is close to the ideal value. The average value of the % elongation [Table 3] from the samples of the deformation defect is 19.04%, which is on the higher side.

The average value of the % elongation [Table 3] from the samples of the deformation defect is 19.04%, which is on the higher side. Again, these shows and increase in the ductility of the samples.From the above test results we can clearly see that the average value of data gathered conforms closely to ideal properties of Ti64, although there was slight deviation in the properties of few samples. Some samples yielded better results to the tune of 11% more than the properties of ideal substrate. This can be attributed to the fact that compressive stresses were induced in few damages. While creating the damage for the deformation defect a force of 3500 N was used to induce the damage to the substrate. The imposition of stress by an external agent usually creates some strain (deformation) in the material, even if it is too small to be detected.

The imposition of stress by an external agent usually creates some strain (deformation) in the material, even if it is too small to be detected. In a solid material, such strain will in turn generate an internal elastic stress, analogous to the reaction force of a stretched spring tending to restore the material to its original undeformed state. This occurs largely due to the transformation of the beta structure to alpha-beta structure which leads to increased strength. Few of the samples also resulted in values lower than that of the ideal samples. This can be concluded from the fact that, few samples had higher heat affected zone during the Laser deposition process. It has been seen that properties of Titanium alter in the areas of the heat affected zone immediately adjacent to the fission zone; specifically those regions which reach temperatures above 1800 F [4]. The percentage elongation changes directly indicate to the change in the ductility of the samples.

The percentage elongation changes directly indicate to the change in the ductility of the samples. The ductility of Titanium drastically changes with its microstructure and the grain size change [5]. From the results its evident that the most fluctuations occurred in the percentage elongation properties due to the change in the microstructure and the grain size of titanium. The results from the UTS, YS tests follow a close proximity to the ideal value of a virgin substrate, referring that the mechanical properties of the repaired samples has not been altered to a large extent.Other factors which generally play a role in determining the properties of Ti64 are the presence of oxygen, nitrogen, hydrogen and carbon, which are interstitial elements, and tends to increase the strength and decrease the ductility. Aluminum is the most important substitutional solid solution strengthener. Its effect on the strength is linear.

Aluminum is the most important substitutional solid solution strengthener. Its effect on the strength is linear. Other, less important sources of strengthening are interstitial solid solution strengthening, grain size effect, second phase (beta) effects, ordering the alpha, age hardening, and the effects of crystallographic texture [2].5. CONCLUSION The properties of the material were enhanced by the compressive stress induced during deformation. An enhancement of the properties can be seen from the test specimens tested from the cutter pull out defect. The properties of the specimens tested with an undercut defect conform to the properties of the ideal condition of substrate. The study was performed on parts supplied by Missouri University of Science and Technology and GKN Aerospace. Anextensive study on the properties of the repaired parts was performed.

The study was performed on parts supplied by Missouri University of Science and Technology and GKN Aerospace. Anextensive study on the properties of the repaired parts was performed. The study included preparing the defect for laser deposition, depositing the Ti-6Al-4V by laser, machining the deposition to complete the repair, and making samples out of the repaired part for property testing. The testing was performed to evaluate the yield strength, ultimate tensile strength, and % elongation of the sample created from the repaired part. The study the current technique of repair as effective.revealed encouraging results, backing6. ACKNOWLEDGEMENT This research was supported by GNK Aerospace and National Science Foundation Grant IIP-1046492, Intelligent Systems Center and Material Research Center at Missouri University of Science and Technology. Their support is greatly appreciated.

Their support is greatly appreciated. The author would like to thanks Todd Sparks from Dept of Mechanical & Aerospace Engineering, Missouri S&T for his continuous support to the research. The author would also like to thanks Laura Bartlett of Dept. of Material Science, Missouri S&T for helping us conduct the tensile strength test, Bob Hribar from machine shop and John Tyler from Rock Mechanics, Missouri S&T for their valuable contribution to the project. Also all people attached directly or indirectly to the project.7. REFERENCES [1] Benjamin, G., Gumenyuk, A.,Rethmeier, M., ., “Laser metal deposition as repair technology for stainless steel and titanium alloys”, Physics Procedia, Vol 39,2012, Pg 376-381[2] Donachie.Jr., Mathew.J., “Titanium – A technicalguide(2nd edition)”, ASM International[3] Potter, Merle C., “Schaum’s outline of dtrength ofmaterial” , McGraw-Hill USA[4] Gordine. J., “weld heat affected zone in titanium” WRCbulletin No 187[5] Zherebtsov.

ditive Manufacturing of 1018 Steel: Process Observations and CalculationsC. M. Knapp * †, T. J. Lienert†, C. Chen†, and D. Kovar*, Solid Free-Form Fabrication Conference: Austin, Texas: August 4-6th 2014Department of Mechanical Engineering, The University of Texas at Austin: Austin, TX 78731 † Material Science and Technology, Los Alamos National Laboratory: Los Alamos, NM 87545ABSTRACTThe temperature distribution in the vicinity of the laser used in direct metal deposition (DMD) plays a critical role in determining the final microstructure and mechanical properties of the deposit and the heat-affected zone (HAZ) within the substrate. Samples were prepared using Laser Engineered Net Shaping (LENSTM) by depositing AISI 1018 steel powder onto AISI 1018 steel substrates in multiple, overwritten passes. The laser power and speed were varied to control the heat input and the rate of cooling.

The laser power and speed were varied to control the heat input and the rate of cooling. The process characteristics were then quantified and compared across the samples to determine the effect of input parameters on the resulting deposit microstructures.INTRODUCTIONAdditive manufacturing has become mainstream in near-net shape fabrication and manufacturing. More recent efforts have been made to enable direct writing of metallic components that can be used in load-bearing components [1]. To realize this objective it is necessary to understand the relationship between the processing parameters, the morphology of the deposit, and micro- and macro-scale defects that can arise from process instabilities.Unlike polymer additive manufacturing routes, where extrusion technologies can be used, direct metal deposition requires melting of powders. There are two common methods to generate the necessary energy densities to melt powders in high fusion temperature metals.

There are two common methods to generate the necessary energy densities to melt powders in high fusion temperature metals. One option is to use an electron beam, which requires that the entire sample be placed under vacuum. However, a more common approach is to use a continuous-wave laser, such as that used in Laser Engineered Net Shaping (LENS™), which is the subject of this research. This research studies the influence of process input parameters, including laser power, laser speed and powder feed rates, on process efficiencies and the morphology of the deposit. We fabricate a sample that consists of fifty overwritten, one-dimensional deposition passes, to produce a 1.5 dimensional “wall” with an intended final height of 2.54 cm.

We fabricate a sample that consists of fifty overwritten, one-dimensional deposition passes, to produce a 1.5 dimensional “wall” with an intended final height of 2.54 cm. Following deposition, the samples were examined for defects and the morphology of the deposit was correlated to processing parameters.EXPERIMENTAL PROCEDUREDEPOSITION SYSTEMA LENSTM MR-7 (Optomec Inc., Albuquerque, NM) was used to produce the samples used in this study. A schematic of this system is shown below. The system consists of acontinuous wave (CW) 1 kW Yb-fiber laser (1070 nm), an aerosol powder feeder, and a deposition head. The location of the deposition head is controlled using three-axis, computer- numerical control.The powder feeding system in the LENSTM MR-7 operates by the rotation of a screw type mechanism, which builds up mounds of powder as it rotates. Argon gas is flowed over the top of the mounds of powder that are built up and caries the powder particles towards the deposition head.

Argon gas is flowed over the top of the mounds of powder that are built up and caries the powder particles towards the deposition head. The powder is fed through four nozzles symmetrically positioned on the deposition head into the focus of the laser beam. As the deposition head moves, molten material is deposited onto the substrate and subsequently cools (shown in orange in Fig. 1).FIGURE 1:SCHEMATIC OF THE LENSTM MR-7 SYSTEM (DEPOSITED MATERIAL SHOWN IN ORANGE) ADAPTED FROM LIOU ET AL. [2]MATERIAL SYSTEMAISI 1018 steel was chosen as the material system to be studied because of its wide availability in plate and powder form, and it affords a relatively straightforward temperature- dependent microstructure from which post-mortem analysis of the thermal characterization can be inferred.POWDEROptomec recommends -80 +325 mesh spherical powder for use in their powder feeders [3]. This size distribution allows for uniform powder supply to the deposition head.

This size distribution allows for uniform powder supply to the deposition head. Powder sizewas determined by mounting powder in epoxy, lightly polishing, and etching to visualize the shape and morphology. The spherical shape and fine grain size shown in Fig. 2 are consistent with the gas atomization process by which the powder was produced by Nuclear Metals Incorporated (Concord, MA, USA).40N e l p m a s n i y c n e u q e r F353025201510500 85 80 95 90 0 15 5 0 0 1 3 1 1 1 Diameter (μm)0 1 10 2 15 2 15 3 10 4 15 4 1FIGURE 2: (LEFT) EXAMPLE 10× MICROGRAPH OF MOUNTED, UNPROCESSED 1018 POWDER. (RIGHT) POWDER SIZE DISTRIBUTION HISTOGRAMA series of micrographs similar to the one shown in Figure 2 were compiled and analyzed using ImageJ (1.47v, NIH, Bethesda, MD, USA) image analysis software. A bias is introduced by the optical measurement of particle size distribution from a random cross section.

A bias is introduced by the optical measurement of particle size distribution from a random cross section. This bias was corrected using the approach presented in Heilbronner et al. [4], to obtain the true powder statistics. The true powder diameter mean μ(cid:2930) is 121 μm and the true standard deviation σ(cid:2930) is 17.3 μm. While this is 20% larger than the ideal diameter recommended by Optomec Inc. of ~100μm, it is well within the operation window when the powder is spherical in nature and flows well.SUBSTRATEAISI 1018 6.35 mm (¼ inch) thick, cold finished, steel substrates were fabricated and marked for three depositions per substrate. The substrates were then bead blasted to decrease reflectivity, and enhance the process efficiency. Four thermocouples each were then spot welded onto the substrate at varying distances from the deposits so that in situ thermal measurements could be taken.

Four thermocouples each were then spot welded onto the substrate at varying distances from the deposits so that in situ thermal measurements could be taken. However the temperature data taken from the thermocouples is outside the scope of this research paper, which focuses on process analysis.PROCESS PARAMETERSIn order to accurately characterize the process, the initial conditions of the depositions, such as laser power and powder feed rate must be experimentally determined. The powder feed rate is set and controlled by a tachometer that measures the rotational speed of the motor used to feed powder. In order to determine the relationship between rotational speed and powder feed rate, three tests were performed by flowing powder into plastic bags for one minute and then weighing the collected material to determine the amount of powder that had been fed in the allotted time.

The three mass flow measurements at varying powder feed revolutions per minute0 5 1yielded a relationship between powder feeder speed and mass flow rate. These points were then used to create a regression analysis to obtain mass flow rate of powder as a function of powder feeder revolutions per minute.To quantify the losses in the system up to the point of the impact of the laser on the substrate, the true laser power was measured using a lollipop-type calorimeter that was placed in the laser focus for a period of 30 seconds. These measurements were done in successive trials to obtain a curve of the actual power produced as a function of laser power demand. The curve was then fit using linear regression so that the actual power could be determined from the power demand. Three trials were conducted with the processing conditions shown in Table 1.

The curve was then fit using linear regression so that the actual power could be determined from the power demand. Three trials were conducted with the processing conditions shown in Table 1. The three trials were conducted under conditions of 1) high powder feed rate, and moderate laser power and laser speed, 2) low powder feed rate, laser power, and laser speed, and 3) high powder feed rate, laser power, and laser speed. A schematic illustration of the sample geometry is shown in Figure 3; a total of fifty successive deposits were overlaid to create a wall that was nominally 2.54 cm (1 inch) tall and 7.62 cm long.TABLE 1: INITIAL CONITIONSTest #1 2 3Initial Powder Feed Rate [g/min]14 3.8 15Laser Power [Watts] 430 290 600Laser Speed [mm/s]6.35 2.11 19.0FIGURE 3: 50 PASS WALL DEPOSITION SCHEMATICPrior to deposition, the part trajectories must be programmed and the specific powder gas flow levels and motor RPM must be set.

The user must then wait for the flow to stabilize before the deposition can be executed. Deposition begins by turning the laser, and loitering at the initial point for fractions of a second prior to the movement of the deposition head. The loitering of the laser is a key difference between selective laser sintering (SLS) and LENSTM. In LENSTM full melting is induced and the powder is deposited into the melt pool [5]. Once the melt pool is initiated the deposition head begins to move and the part is slowly built up, pass-by-pass. The optical setup used to produce the deposits maintained the focus of the laser ~300μm beneath the surface of the layer to be deposited. After the completing the first pass, the laser was broughtback to the initial position. The successive pass was made in the same way except that the height of the focus was increased by 0.508 mm.

After the completing the first pass, the laser was broughtback to the initial position. The successive pass was made in the same way except that the height of the focus was increased by 0.508 mm. The resultant component is shown in Figure 4.FIGURE 4: COMPLETED DEPOSITIONS (NOTE ASYMMETRIC OXIDE FORMATION)ANALYSIS OF DEPOSITSAfter deposition, the samples were coarsely cut from the plate using an abrasive cut-off saw. Regions of interest were further sectioned to smaller sizes using a low speed saw to minimize the damage to the specimen. After cutting and mounting in Buehler Kwick Kit long cure epoxy (Buehler, Lake Bluff, Illinois, USA), the specimens were ground and polished. The samples were ground using SiC paper and then polished using diamond suspensions on an automatic polisher (Struers Rotopol 11 and RotoForce 1, Westlake, Ohio, USA). A sequence of suspensions starting with 9μm, 3μm, and 1μm was used.

A sequence of suspensions starting with 9μm, 3μm, and 1μm was used. The force applied during polishing was 40 N and each of the polishing steps was conducted for 4 minutes. After polishing with the diamond suspensions, a final polishing was conducted with 0.05μm alumina suspension at a force of 30 N for 3 minutes. The samples were then etched using 5% nitric acid and 95% ethanol (5% Nital) to reveal the microstructure. Stereo microscopic images of section of the samples are shown in Fig. 5.FIGURE 5: STEREO-MICROSCOPE CROSS SECTION IMAGES OF THE ONE INCH WALL DEPOSITION FROM TESTS 1(A), 2(B)(SECTIONED TO FIT IN MOUNT), AND 3(C)Note that sample 2 (Fig.

5.FIGURE 5: STEREO-MICROSCOPE CROSS SECTION IMAGES OF THE ONE INCH WALL DEPOSITION FROM TESTS 1(A), 2(B)(SECTIONED TO FIT IN MOUNT), AND 3(C)Note that sample 2 (Fig. 5b) was significantly taller than the other two samples; this sample was cut at roughly the mid-plane and the top half of the sample was placed next to the bottom half in order to fit the sample in the mount used for polishingPROCESS MEASUREMENTSIn order to compare input parameters across tests, specific energy was used and defined to be the supplied laser power divided by the traverse speed. High specific energy occurs at high power and low speed, and the low specific energy occurs at low power and high speed. The specific energy was plotted with the powder feed rates that were supplied to the deposition to show the relationships between laser power, speed, and powder feed rate (Fig 6).

The specific energy was plotted with the powder feed rates that were supplied to the deposition to show the relationships between laser power, speed, and powder feed rate (Fig 6). However, observation of the deposits revealed that certain powder feed rates were not conducive to stable deposition (shown in red) while others were stable (show in green).30] n i m / g [2520StableUnstablee t a R w o l F r e d w o P1510500100 50 Specific Energy [kJ/m]150FIGURE 6: POWDER USE EFFICIENCY AND DEPOSITION STABILITY AS A FUNCTION OF SPECIFIC ENERGYStability in this study refers to how closely the actual deposition height corresponds to the z-step of the deposition head. If the deposition height is not high enough to accommodate the height of the programmed z-step, the position of the laser focus will begin to rise so that it is no longer buried in the substrate.

If the deposition height is not high enough to accommodate the height of the programmed z-step, the position of the laser focus will begin to rise so that it is no longer buried in the substrate. This is a result of not enough powder being fed into the molten pool and is referred to as an under built condition. However, if too much powder is fed, the deposit will get too close to the laser and change the laser focus condition. This condition results in inconsistent layer deposits and is referred to as an over build condition. While the overbuild condition occurred to a small degree in test 2, for most of the trials, the under build regime dominated and was constantly requiring corrections to the supplied powder flow rate.OBSERVATIONSThe effects of process variation across test settings are immediately apparent by examining of the metallographic cross-section. Shown in Figure 7(a) is the difference in deposit height between tests 1 and 2.

Shown in Figure 7(a) is the difference in deposit height between tests 1 and 2. Test 1 has the proper balance between powder feed, laser speed, and laser power; resulting in an individual layer height corresponding to the z-step of the deposition head. This balance results in a constant distance between the deposition head and the top of the deposit. The constant distance allows for a constant melt pool size, consistent width of deposited material, and a final component that accurately reflects the dimensions of computer model. Conversely in test 3 in Figure 7(a) the height is ~70% of the intended 1 inch. The initial settings for this deposit had a powder supply deficit relative to laser power and speed, which led to a layer height that was insufficient to match the z-step of the deposition head.

The initial settings for this deposit had a powder supply deficit relative to laser power and speed, which led to a layer height that was insufficient to match the z-step of the deposition head. The deficient layer height compounds over successive passes, yielding a component that is significantly under built.FIGURE 7: (LEFT) STEREOMICROGRAPH OF GEOMETRIC VARIATION BETWEEN TESTS 1 AND 3. (RIGHT) LARGE INTERNAL PORE DEFECT IN TEST 3.Additionally in Figure 7(b) shows the macroscopic defects in the build that can form as a result of unstable deposition. The width of deposit in test 3 in nominally 1.4 mm, yet as a result of unstable deposition there is a 260 μm diameter closed pore in the center, with other porosity around it.

The width of deposit in test 3 in nominally 1.4 mm, yet as a result of unstable deposition there is a 260 μm diameter closed pore in the center, with other porosity around it. These defects are not visible from the surface but would severely degrade performance if this was used as a structural component.MELT POOL INSTABILITYThe recommended operating alignment for the LENSTM MR-7 is a focus that lies 0.305 mm (0.12 in.) above the deposit plane, as shown in Fig. 8. Note that if the substrate is pulled away from this point so that the spot size is increased, an overfeed situation occurs and the resulting deposit thickness is greater than what is intended. Conversely, a smaller spot size results in an underfeed situation and the deposit thickness is less than anticipated.

Conversely, a smaller spot size results in an underfeed situation and the deposit thickness is less than anticipated. Also note that these effects are magnified when multi-pass depositions are conducted.FIGURE 8: POSITIVE DIVERGENCE FEEDBACKFrom the instabilities that were observed during deposition with a positive divergence feedback, it is clear that this is not the operational regime desired for such a dynamic process. The ideal set of process parameters result in a process that self-corrects so that it does not become unstable. One way to create a passive negative deposition feedback, is to move the deposition location to the far side of sharp laser focus. When the laser and deposition head are aligned to create this condition it will result in the schematic shown in Figure 9.FIGURE 9: NEGATIVE DIVERGENCE FEEDBACKWhen operating under positive divergence feedback, the deposition will self correct.

For example, if the process is operating in an over feed condition; the spot size is decreased as is the melting efficiency and layer height. In contrast, if the process is operating in an under feed condition, the spot size is increased, as is the melting efficiency and layer height. Thus, multi- pass depositions will auto-correct over sequential layers.CONCLUSIONSInitial alignment and input parameters have a significant impact on the depositions produced by the LENS process. Deposition instability results from an inequality between laser speed, power, and powder feed, as well as a default positive divergence feedback. This instability leads to significant deviations between the intended height of a multi-pass deposit and the actual height.

This instability leads to significant deviations between the intended height of a multi-pass deposit and the actual height. While inequalities persist regardless of feedback conditions, it was suggested that the divergence between intended and actual height can be eliminated by refocusing the laser so that deposition occurs on the opposite side of sharp laser focus (e.g. above the surface of the preceding deposit). Future research will explore the effects of operating on the far side of sharp focus, and further development of analytical models for stable deposition.ACKNOWLEDGEMENTSThe authors would like to acknowledge the funding for this research by Los Alamos National Laboratory’s Material Science and Technology Division, in collaboration with the University of Texas of Austin.REFERENCES[1][2][3] [4]Lewis, Gary K., et al. Directed Light Fabrication of Near-Net Shape Metal Components. s.l.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAdditive Manufacturing of Alumina Components by Extrusion of in-situ UV-cured PastesLok-kun Tsui*, Erin Maines†, Lindsey Evans†, David Keicher@, and Judith Lavin†Center for MicroEngineered Materials, University of New Mexico, NM 87106 †Advanced Materials Laboratory, Sandia National Laboratories, NM, 87106 @Integrated Deposition Solutions, Inc. NM, 87110AbstractAdditive manufacturing of ceramic materials is an attractive technique for rapid prototyping of components at small scales and low cost. We have investigated the printing of alumina pastes loaded at 70-81.5 wt% solids in a UV curable resin. These can be deposited by extrusion from a syringe head on a Hyrel System 30M printer. The print head is equipped with an array of UV LEDs, which solidify the paste without the need for any applied heating.

The print head is equipped with an array of UV LEDs, which solidify the paste without the need for any applied heating. Parameters optimized include print speed, layer height, applied force, and deposition rate. Using A15 alumina and submicron A16 powder precursors, we can achieve bulk densities of 91% and 96% of theoretical density respectively. The influence of dispersants and surfactants added to the powder on the rheology of the pastes, the print process parameters, and the quality of the final components are also investigated.IntroductionAdditive manufacturing of ceramic components enables the prototyping of materials and processes at small scales and in geometries difficult to produce with conventional ceramic fabrication techniques [1].

This technique has been used in applications for structural ceramics, biologically compatible components, electronics, and chemical sensors.[1–6] Processes available for additive manufacturing of ceramics include robocasting,[7] where a paste containing ceramic powder and an aqueous or organic binder is extruded from a computer controlled printer onto a heated substrate; freeze form extrusion fabrication, where a paste is extruded onto a cold surface and frozen in place [8]; and stereolithographic printing where a slurry of material is exposed to a 2D pattern of light produced by a projector [4,9,10]. An alternative to these techniques combines the use of light-curable pastes with extrusion by using a UV-light source to cure the pastes as they are being printed. This allows the use of small quantities of material in cases where materials are expensive or scarce and the light source removes the need for a heating or cooling system for solidification. Faes et al.

Faes et al. have investigated the extrusion of UV-curable pastes with zirconia and achieved 92% theoretical density [11,12].We report on the printing of UV-curable pastes containing alumina powders from a commercially available extrusion printer, including the modifications made to the printer in order to monitor the force applied during deposition. Paste composition, printing parameters, and heat treatment conditions were optimized to maximize the final sample density. We are able to achieve up to 90-92% and 94-96% of theoretical alumina bulk density using A15 and A16 alumina951precursors respectively. It is necessary to understand where the defects, which cause deficiencies in final density, are located. Electron microscopy in principle allows the imaging of defects of printed parts, but imaging is limited to only the surface or the sample must be sectioned destructively to evaluate the interior structure.

On the other hand, x-ray computed tomography imaging is a technique used to non-destructively visualize the location and distribution of defects inside these parts.Experimental SectionA16 and A15 alumina powders were obtained from Almatis with the peak of the particle size distribution being 0.5 µm for A16 and a bimodal distribution in A15 with peaks at 0.5 µm and 4 µm. The powders were prepared by ball milling with 5 mm yttria-stabilized zirconia beads in either ethanol with 0.5 wt% oleic acid or toluene with 2 wt% Hypermer KD1 dispersant. After milling, the solvent was allowed to evaporate at room temperature and then the powders dried in a drying oven at 80oC overnight. The dry powder was then crushed with a mortar and pestle and sieved using a No. 40 sieve (425 µm mesh).Pastes were prepared using Tethon3D Genesis UV-curable resin with a Thinky centrifugal planetary mixer.

40 sieve (425 µm mesh).Pastes were prepared using Tethon3D Genesis UV-curable resin with a Thinky centrifugal planetary mixer. For each batch of paste, the alumina was added into the resin in 1/3 size batches and mixed at 2000 RPM for 1 minute at ambient atmospheric pressure. The completed paste was then loaded into a 5 mL syringe and centrifuged at 2500 RPM for 1 minute. Finally, the syringe was placed into the Thinky planetary mixer and mixed again at 2000 RPM under 7 kPa vacuum to degas the paste. Paste loading was varied between 71-82 wt% alumina. A table of pastes evaluated in this set of experiments is shown in Table 1.The printing was performed using a Hyrel System 30M 3D printer which can accept a variety of heads including traditional fused deposition modeling printing of plastic components and extrusion of slurries and pastes from a syringe head.

We have modified the CSD-05 syringe head with a force sensor (Omega LCFL-75 coupled with a DMD4059-DC signal conditioner and amplifier). The signal is received with a National Instruments DAQmx USB-6218 controlled by LabVIEW. A schematic of the force sensing system is shown in Figure 1(a) and a photograph of an unmodified syringe head is shown in Figure 1(b). The Hyrel System 30M may be controlled by standard GCODE generated from 3D printing slicer software such as the included Slic3r or Ultimaker Cura. For the results reported here, we used Ultimaker Cura v. 3.0.2 to generate toolpaths. The CSD-05 head includes three 365 nm UV LEDs which operate at a power density of 0.25-0.5 W/cm2 during printing. Jensen Global opaque Luer lock syringe tips with nozzle diameter 0.025" (635 µm) was used for all printing experiments. Glass microscope slides (75 x 50 x 1 mm) were used as substrates. A list of other printing parameters is shown in Table 2.

Glass microscope slides (75 x 50 x 1 mm) were used as substrates. A list of other printing parameters is shown in Table 2. Simple geometric structures were chosen for testing including a 7.5 x 7.5 x 7.5 mm cube, a 7.5 mm diameter x 7.5 mm height cylinder, and a ring with 8.0 mm outer diameter, 4.8 mm inner diameter, and thickness of 4.6 mm. After the components were printed they were exposed to high intensity UV-LED radiation from a Phoseon UV light source with a power density of 300 W/cm2. Samples underwent burnout and sintering in air and the heat treatment parameters are shown in Table 3. Heat treatment processes were performed in Carbolite CWF 12/5 and Deltech DT-37FL-10-E3504 box furnaces for burnout and sintering respectively.952Figure 1. (a) Schematic of the syringe head modified with a load cell and signal conditioner to perform force measurements as a function of time.

(a) Schematic of the syringe head modified with a load cell and signal conditioner to perform force measurements as a function of time. (b) A photograph of the syringe print head is shown.Paste Name Alumina Type and Loading Oleic Acid Hypermer KD1 Genesis Resin A15G81.5 A16G71 A16G73 A16G77 A16G80 A16G82A15 81.5 wt% A16 71.0 wt% A16 73.0 wt% A16 77.0 wt% A16 80.0 wt% A16 82.0 wt%0.4 wt% 3.6 wt% 3.7 wt% 3.8 wt% 4.0 wt% 4.1 wt%0 wt% 1.4 wt% 1.5 wt% 1.5 wt% 1.6 wt% 1.6 wt%18.1 wt% 24.0 wt% 21.8 wt% 17.7 wt% 14.4 wt% 12.3 wt%Table 1. Paste compositions studied for printing. All values shown are wt% with respect to total paste mass.Parameter Name Extrusion MultiplierA Layer Height Computed Line Width Infill Top and Bottom Layers Printing Substrate Temperature Printing Speed Travel SpeedValue 0.2 200 µm 200 µm 90% Disabled Room Temperature (22oC) 4 mm/s outer perimeter, 7 mm/s infill 17 mm/sTable 2.

Printing Parameters for deposition of alumina pastes with the Hyrel System 30M. AExtrusion multiplier – unitless tuning factor in the Hyrel control software for flow rate.953Temperature (oC)225 625 1000 251625 25Ramp Rate (oC/min) Burnout (Air atmosphere) 0.5 0.5 1.0 -3.0 Sintering (Air atmosphere) 3.0 -5.0Hold Time (Hours)0.5 1.0 1.0 0.02.0 0.0Table 3. Heat treatment parameters for burnout and sintering for alumina components.The as-deposited and post-heat treatment geometry of the samples were measured by optical microscopy using a Keyence Optical Microscope. Density measurements of the components were performed using the Archimedes technique in DI water and compared with the theoretical density of alumina (3.99 g/cm3 [13]). X-ray tomography was obtained on the samples using a North Star Imaging industrial microfocus CT instrument with 150 keV x-rays and imaging pixel resolution of 8.3 µm.

X-ray tomography was obtained on the samples using a North Star Imaging industrial microfocus CT instrument with 150 keV x-rays and imaging pixel resolution of 8.3 µm. Rheological measurements of the paste were performed with a Haake Mars 2 35 mm cone-plate rheometer at 1 degree. An initial constant stress was of 3000 Pa was applied for 120 seconds and the paste was allowed to equilibrate for 500 seconds. A shear rate sweep was performed from 0.11 Hz to 1000 Hz.Results and Discussion1. Paste Rheology and Force MonitoringInitial sets of tests to determine rheological properties, applied force during printing, and print accuracy and force as a function of time were performed on pastes consisting of 81.5 wt% A15 alumina (A15G81.5 in Table 1). Rheological properties of the paste are shown in Figure 2(a) for three sequential runs. The paste exhibit shear thinning behavior, which is desirable for an extrusion application.

Rheological properties of the paste are shown in Figure 2(a) for three sequential runs. The paste exhibit shear thinning behavior, which is desirable for an extrusion application. The paste should be self-supporting when no force is applied and flow easily when sheared in the syringe. We observe hysteresis in the scans with the return scan showing lower viscosity than the onset scan. Figure 2(b) shows the force as a function of time for constant flow of material at a dispensing rate of 1.1 µL/s. In the first 30 seconds after the dispense head contacts the syringe plunger, the force rises linearly until material begins to extrude from the nozzle. The force then reaches a steady state of 45N (200-350 seconds). Dispensing was stopped at 350 seconds and we observe a decay to a resting force of 30N as material continues to be extruded even after motorized dispensing is stopped. This resting force corresponds to the minimum force necessary for the paste to flow out of the syringe nozzle.

This resting force corresponds to the minimum force necessary for the paste to flow out of the syringe nozzle. The addition of the force measurement device to the 3D printer allows real-time monitoring of the applied force during printing to verify whether the material is being consistently extruded. Motor belt skipping can also be detected by watching for sudden decreases in applied force. This occurs when the necessary torque needed to extrude more material exceeds what the stepper motor in the Hyrel CSD-05 head can supply.The print speeds and extrusion factor shown in Table 2 were the result of adjustments so that continuous layers were deposited for the cubes and cylinders. An excessive extrusion factor954or slower print speed results in over-extrusion and loss of resolution, while too small of an extrusion factor or too fast of a print speed results in under extrusion and leaves large voids in the structure.

We determined that the smallest height that we could consistently print structures with was 200 µm, as any smaller would require extrusion rates too which corresponded to applied forces too small to induce material flow from the nozzle.Figure 2. (a) Viscosity vs. shear rate of 81.5 wt% A15 in Genesis showing three sequential scans. Arrows indicate the direction of onset and subsequent return scans. (b) Force curve for this paste with a constant dispensing rate with showing three distinct regions corresponding to onset (1), steady-state (2), and decay to idle force (3) after dispensing is stopped at 350 seconds.2. Print Accuracy3D models and optical microscopy of some simple geometric structures are shown in Figure 3 and a comparison of the expected parameters and the printed parameters are collected in Tables 4 and 5.

The axes conventions used are as follows: Z represents the direction perpendicular to the substrate, while X and Y represent directions in parallel to the substrate. The average sizes of the as deposited structures are typically within 100 µm of the target structure. The main aspects of the printer and print head which impose limitations on resolution are the layer step height of 200 µm and the nozzle diameter of 635 µm. The samples exhibit a higher shrinkage in the Z direction due to the greater density of voids that are eliminated in that direction during sintering. The bottom surface tends to be smooth since it is in contact with the glass substrate during printing while the top surface shows roughness due to the printing of lines which are not spread out against a substrate.

Optical microscopy of post-sintered (top row) Cube and (middle row) cylinder. The images show the samples in different views: (left) Side view. (center) Top view. (right) Bottom view. 3D models of the test structures are shown on the bottom row with the axes convention.9573. Component Density and Internal DefectsThe density of the components created with A15 is only about 90-92% of theoretical alumina density. Per specifications received from the manufacturer, A15 contains a bimodal distribution of particle sizes peaked at 0.5 µm and 4 µm. The larger particles in this powder have lower surface area compared to that of A16, and therefore the driving force for sintering is weaker. To remedy this we moved to using A16 alumina with a unimodal particle size distribution with a peak at 0.5 µm. The smaller particle size produces a mixture with higher viscosity, necessitating the use of dispersants. With only OA, the maximum loading that could be extruded on our printer is 71 wt%.

The smaller particle size produces a mixture with higher viscosity, necessitating the use of dispersants. With only OA, the maximum loading that could be extruded on our printer is 71 wt%. The introduction of Hypermer KD1 in place of oleic acid at loadings of 2 wt% KD1 resulted in a mixture with a consistency of water up to 82% where a transition to a shear thickening material occurs. This behavior was observed qualitatively by disturbing the paste with a wooden spatula. In the presence of both 5 wt% OA and 2 wt% KD1 the paste is a shear thinning material, which can be extruded at loadings up to 82 wt% (Pastes A16G71 to A16G82 in Table 1). The oleic acid acts as a lubricant so that the particles can slide past each other. Otherwise, the particles jam together during the application of a shear force, resulting in shear thickening behavior.Experiments conducted using A16 used the 8.0 mm diameter ring structures to determine shrinkage and theoretical density.

Components printed using 71 wt% A16, without accounting for shrinkage, exhibited a maximum theoretical density of 96%. The next set of experiments investigated the influence of powder loading between 71 and 82wt% on the density and shrinkage of the samples. Microscopy of a ring structure printed with 82 wt% A16 is shown in Figure 4. In this set of samples, the components were printed with a linear scaling factor of 1.26 to 1.39x to compensate for shrinkage during sintering based on the expected volumetric shrinkage from the loading of the powders. Figure 5 shows trends in the final post-sintered density as a function of sample loading of the components determined using the Archimedes technique. In the range of loadings tested, the density increases from 88% to 95% of theoretical density.

In the range of loadings tested, the density increases from 88% to 95% of theoretical density. The reason that the 71 wt% sample has a lower density than the initial test, which achieved 96%, is that the larger components have been scaled for shrinkage requiring more layers, and therefore, contain a greater number of defects. We also observe a decrease in the fraction of open pores with higher loadings— this is due to less surface cracking when less shrinkage occurs in higher loading samples. An increase in the fraction of closed pores is observed with increasing loading. This may be a result of the high viscosity of the paste preventing it from easily flowing into and filling internal voids. Figure 5(c) shows trends in the shrinkage as the alumina solids loading is increased. In all cases, the shrinkage in the Z-thickness of the rings is greater than the thickness in the X-Y direction, indicating that shrinkage compensation must be greater in the Z direction than the X-Y direction.

The shrinkage for thickness, outer diameter, and volumetric shrinkage is observed to decrease linearly with increasing alumina solids loading.958Figure 4. Side view and top view of an alumina ring structure printed with A16 alumina after sintering.Figure 5. (a) Fraction of theoretical density as a function of alumina loading shows an increase in the density of the components. (b) The fraction of open pores decreases as a function of alumina loading while the fraction of closed pores increases. (c) Shrinkage as a function of alumina solids loading decreases linearly in for both length parameters and volumetric shrinkage. Error bars represent standard deviation of 6 samples.9594. X-ray Computed TomographyFinally, X-ray tomographic imaging is used to non-destructively probe for detects in the interior of additively manufactured ceramics. Cross sections in the X-Y and X-Z planes of the ceramic cylinders are shown in Figure 6.

Cross sections in the X-Y and X-Z planes of the ceramic cylinders are shown in Figure 6. Dark areas represent voids, while light areas represent solid material. The x-ray imaging indicates small cracks embedded inside the alumina structure, which are responsible for the deficiency in density. In the X-Y plane, defects are located close to the outer edges of the ring structure. In the X-Z plane, the defects are located in lines parallel to the plane of the layers during printing. This indicates that one source of defects is the deposition of layers on top of each other inducing the formation of voids at the interface. Careful tuning of the deposition parameters will be needed in order to minimize void formation between the layers.Figure 6. X-ray computed tomography of an A16 ring structure.

Careful tuning of the deposition parameters will be needed in order to minimize void formation between the layers.Figure 6. X-ray computed tomography of an A16 ring structure. (a) Cross section through a ring structure in the X-Y plane and (b) cross section through a ring structure in the X-Z plane.ConclusionsWe have studied the use of UV-curable alumina pastes to additively manufacture ceramiccomponents. The rheology and force as a function of time for a constant dispensing rate of the pastes were examined. The geometry and final sintered density of the printed components was measured, and higher densities were observed by using smaller particle size alumina precursors allowing us to reach a maximum density of 95-96% of the theoretical density of alumina.

X-ray tomography was then used to visualize the location and distribution of defects which contribute to deficiencies in component density.960AcknowledgementsWe acknowledge Nelson Bell (Sandia National Laboratories) for assistance in obtaining and analyzing rheological data. We thank Joseph A. Romero (Sandia National Laboratories) for assistance in obtaining X-ray tomography measurements on our samples.This work was funded by the Sandia National Laboratories LDRD program. Sandia National Laboratories is a multi-mission laboratory managed and operated by National Technology & Engineering Solutions of Sandia, LLC, a wholly owned subsidiary of Honeywell International Inc. for the US Department of Energy’s National Nuclear Security Administration under contract DE- NA0003525.References[1] N. Travitzky, A. Bonet, B. Dermeik, T. Fey, I. Filbert-Demut, L. Schlier, T. Schlordt, P.[2][3] [4][5] [6][7] [8] [9]Greil, Adv. Eng. Mater. 16 (2014) 729–754. P. Sarobol, A.

After creating carbon polymer composite filaments, using a commercial 3D printer, samples were printed and tested for mechanical and electrical properties. The measured values for these composites were compared to those obtained for pure ABS and pure PLA. It was found that by using only 2% SCF it was possible to achieve a 22% increase in tensile strength with no significant impact on printability. With addition of graphene, PLA was made to be conductive. These results show the feasibility of developing new materials for 3D printing that will create structurally sound and conductive designs.IntroductionAdditive manufacturing processes have the potential to alter the engineering landscape. The untapped resource of carbon polymer composite printing has the ability to create lightweight components of high strength, taking advantage of the repeatability and intricacy of additive manufacturing.

It is also possible to use these materials to create electrical components due to the electrical properties of some carbon polymer composites.The carbon structures used throughout this investigation were graphene and short carbon fibres. Graphene is an incredibly strong material with a tensile strength of 130000 MPa (Potts et al. 2011) over 70 times greater than that of steel. It is composed of a single layer of carbon atoms formed inside a honeycomb lattice (Akhaven. 2009). Graphene also provides a high conductivity, measured up to 3500S/m in stacked graphene oxide (Potts et al. 2011). Carbon fibre is currently used widely in applications demanding high strength and low weight, such as aerospace, high performance marine, automotive and sporting equipment (Fu et al. 2000). Composed of interlocking sheets of graphite (MIT, 2015), it has a tensile strength greater than 3GPa (Chen et al. 2004), and low electrical resistivity of 10-2-10-3 Ωcm (Chen et al.

2000). Composed of interlocking sheets of graphite (MIT, 2015), it has a tensile strength greater than 3GPa (Chen et al. 2004), and low electrical resistivity of 10-2-10-3 Ωcm (Chen et al. 2004).It is well known that the addition of a small amount of carbon nanostructures within a polymer matrix can lead to enhanced material properties. Kuilla et al. (2010) find that by adding 0.1% by weight of graphene to an epoxy matrix improved flexural strength by 87%. Similarly, addition of 0.1% by volume of graphene sheets to a polystyrene matrix increased electrical conductivity by a factor of 4 x 1014. Cong et al. (2015) found a 21% increase in 864ultimate tensile stress with the addition of 7.5% short carbon fibers in an ABS polymer matrix. In this work these carbon polymer composites with enhanced material properties were combined with the 3D printing process.

In this work these carbon polymer composites with enhanced material properties were combined with the 3D printing process. This was done to examine the extent to which these material properties were conserved through the 3D printing process and the viability of this production technique.Experimental TechniquesFigure 1- Schematic of specimens (dimensions in mm)The investigation involved analysis of structural and electrical properties of 3D printed carbon polymer composite components. To achieve this, short carbon fibres of 212µm were mixed with commercially available ABS pellets in a single screw Filabot filament extruder in specified ratios. These filaments were then extruded at approximately 178oC and printed into tensile test coupons. The 3D printer used was a commercially available Leapfrog 3D Creatr. The printing parameters used were a print temperature of 230oC with a bed temperature of 80 oC.

The 3D printer used was a commercially available Leapfrog 3D Creatr. The printing parameters used were a print temperature of 230oC with a bed temperature of 80 oC. Test coupons were then analysed for conductivity using a high accuracy ohm meter and in tension until failure using a MTS Criterion loading frame. Pure ABS, 1wt% SCF/ABS, 2wt% SCF/ABS, PLA and PLA/Graphene coupons were subject to this test regime. The PLA/Graphene filament was sourced from Black Magic 3D (2015).ResultsThe tests show that it is possible to use the enhanced properties of carbon polymer composites in the 3D printing process. Figure 2 shows the stress strain curve of ABS printed samples against the curves of 1wt% and 2wt% SCF composite printed samples. Table 1 compiles these results showing that a strength increase of 22% was observed with only 2wt% addition of SCF to the polymer matrix.

Table 1 compiles these results showing that a strength increase of 22% was observed with only 2wt% addition of SCF to the polymer matrix. Whilst this increase in tensile strength is significant, Figure 2 also shows that whilst ABS exhibits plastic yielding, the SCF composite specimens do not behave in the same manner and their ductility is reduced.Figure 3 shows the stress strain interaction of the PLA samples against those containing PLA/Graphene. It can be seen that despite the superior properties of graphene, the printed samples do not display the enhanced material properties, with these samples having a 54% lower tensile strength.

Using density analysis, it was found that a 20% infill lead to the printed cross section being 69% of the total area. Using this analysis, the material properties were calculated and are listed in Table 1. As shown by Table 1, even with theoretical scaling, the PLA coupons do not achieve the usual tensile strength of 56.6MPa (Tymerak et al. 2014) whilst the ABS coupons also do not reach the usual value of 42.5MPa (Teststandards. 2015). This may be due to a lack of bonding of layers within the 3D printing process, and delamination occurring between the layers of printed material.The addition of carbon structures to a polymer matrix did affect the conductivity of the materials. As seen in Table 2, addition of SCF did not change the conductivity of the ABS drastically, whilst the graphene samples provided conductivities five orders of magnitude higher than the SCF specimens. A reading could not be obtained for the PLA samples due to the apparent insulating property of the material.

A reading could not be obtained for the PLA samples due to the apparent insulating property of the material. However through the 3D printing process, the conductivity of the sample was reduced as shown through the pure filament having a conductivity over three times higher than that of the printed PLA/Graphene specimen.Table 2 – Conductivity of Tested samplesSpecimen Pure PLA Coupon PLA-graphene Coupon PLA – graphene Filament Pure ABS Coupon 1wt% SCF-ABS Coupon 2wt%SCF-ABS CouponConductivity (S/m) N/A 11.24416 38.75969 3.39E-05 3.80E-05 3.82E-05One of the largest factors that influences the results obtained is the printability of filaments. Using the Leapfrog Creatr 3D in this investigation, the print head diameter was 0.5mm. Throughout the printing process, the clogging of this nozzle was the most significant hindrance of results, causing many prints to be incomplete.

Throughout the printing process, the clogging of this nozzle was the most significant hindrance of results, causing many prints to be incomplete. Whilst higher compositions of carbon structures were intended to be explored, the 2wt% SCF formed the highest printable filament and hence shows a limitation of this investigation.867DiscussionIt is not surprising that an introduction of SCF into an ABS matrix increases tensile stresses, confirming previous results from Cong. Et al (2015). However this investigation shows that a significant increase in tensile strength of 22% is achievable with a low percentage of SCF and no impact on printability. Whilst Cong. et al (2015) were able to achieve a 21% increase in their printed samples with 7.5% SCF introduction, this investigation shows that it is possible to achieve a greater increase of 22% with only a 2wt% addition.

This may be attributed to the fact that Cong et al (2015) used 150µm long SCF whilst this investigation used 212 long µm fibres. Cong et al. (2015) also find that increasing the fibre length from 100 to 150µm increases tensile strength by close to 10%, which is consistent with our result.Figure 4 – Image of failed Graphene – PLA specimensThe introduction of the carbon structures into the polymer matrix also altered the failure mode. The ABS samples failed after a large amount of plastic deformation was seen by the material, whilst the SCF and graphene composite samples exhibited brittle characteristics as shown in Figure 3. Figure 4 shows that the failure did not exhibit signs of yielding, suggesting a brittle failure mode. Similarly It was observed on most samples that the failure occurred along a 45 degree axis. This is largely due to the structure of the 3D printed sample, determined by the pattern of infill set up in the gcode of the 3D printer process.

This is largely due to the structure of the 3D printed sample, determined by the pattern of infill set up in the gcode of the 3D printer process. Determined by the gcode, for all infill, material was laid down by the printer head at an angle of 45 degrees to the coupon’s longitudinal axis. This explains the 45 degree failure observed in most sample’s infill, as the strands of material did not have to break, but rather, the bond between each strand of printed material failed first. This shows that it is not entirely the material failing, but rather a structural flaw in the design of the coupon leading to premature failure.However, it may not have been the carbon structures themselves that lead to brittle failure, as the tensile strength for SCF and graphene are 2950 MPa (Fu et al. 2000) and 130000MPa (Potts et al. 2011) respectively, and it is unlikely that these values were reached during the testing process.

2000) and 130000MPa (Potts et al. 2011) respectively, and it is unlikely that these values were reached during the testing process. Instead however it is more likely that the bonds between the polymer and the carbon structures within the matrix fail with brittle characteristics. This effect may be reduced with an increase in length of the carbon structures. Potts et al. (2011) sight this effect, stating that as a result of scanning electron microscopy study, two868dimensional stiffness of graphene is unable to be utilised by three dimensional polymer sample. To reduce the influence of this effect, Mark Forged (2015) increased length of printed sampled by developing a process to lay continuous strands of carbon fibres within a polymer matrix. The company claims strength increase up to 27 times compared with a pure ABS print. However, this process requires specialised equipment and does not follow the standard FDM printing process.Both SCF and graphene are highly conductive.

However, this process requires specialised equipment and does not follow the standard FDM printing process.Both SCF and graphene are highly conductive. However this conductivity is not reflected in the printed composite materials. Whilst carbon fibres have a low electrical resistivity of 100-1000 S/m (Chen et al. 2004) and graphene has a high conductivity of up to 3500S/m (Potts et al. 2011), as shown by Table 2, the conductivities of the SCF composite samples were near zero rendering them poor conductors, and the conductivity of the graphene filament was roughly 10 times lower than the value measured by Potts et al. (2011). A low concentration of the carbon structures within the sample provides one possible explanation for the lack of conductivity observed in the samples. The printing process also significantly altered the conductivity of samples. As shown in Table 2, the PLA/graphene filament provided conductivity over three times greater than the printed sample.

The printing process also significantly altered the conductivity of samples. As shown in Table 2, the PLA/graphene filament provided conductivity over three times greater than the printed sample. Inherent in the FDM printing process is a deposition of very fine material that is often laid in an orientation not following the path of current flow. Similarly, between layers there may not be a path for electrons to flow due to small air voids or a lack of contact of conductive carbon structures within the sample. These variables significantly alter the conductivity of samples through the printing process when compared to the corresponding filament.ConclusionIt can be seen from the investigation that it is possible to enhance the mechanical and conductive properties of 3D printed structures with the addition of carbon structures within the polymer matrix. The conductivity of a pure PLA sample was increased with the introduction of graphene into the polymer.

The conductivity of a pure PLA sample was increased with the introduction of graphene into the polymer. Hence the 3D printing process involving the addition of graphene to polymer filaments was effective in increasing the conductivity of printed structures. This could lead to potential application of this process for 3D printed electrically conductive components such as circuit boards. The 3D printed coupons comprised of PLA-graphene composite materials exhibited reduced tensile strength and stiffness characteristics. These structures also exhibited brittle failure and no yielding compared with the ductile PLA coupons. Hence the 3D printing process involving graphene additives was ineffective in enhancing the material properties of polymer structures. The investigation showed that it was possible to enhance the tensile strength of ABS 3D printed structures with the addition of SCF.

The investigation showed that it was possible to enhance the tensile strength of ABS 3D printed structures with the addition of SCF. 2wt% SCF-ABS 3D printed structures exhibited higher tensile strengths, however showed less ductility than pure polymer structures. These enhanced properties show potential application of this process in creating stronger 3D printed components for high accuracy applications such as aerospace and industry. The conductivity of ABS filament was not increased to highly conductive levels by the introduction of SCF filaments. The addition of the SCF at the 2wt% level was unable to create viable conductivity within the sample for applications requiring electrical conductivity.869ReferencesAkhavan O, 2009 ‘The effect of heat treatment on formation of graphene thin films from graphene oxode nanosheets’ Carbon, Feb 2010, vol.48, no.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAdditive Manufacturing of Cu on 316L Stainless Steel via Inconel 718 Intermediate LayersXinchang Zhang 1, 2, *, Tan Pan 2, Yitao Chen 2, Frank Liou 21 Idaho National Laboratory, Idaho Falls, Idaho, 83401, USA 2 Department of Mechanical & Aerospace Engineering, Missouri University of Science and Technology, Rolla, Missouri, 65409, USA * Corresponding author: xinchang.zhang@inl.govAbstractJoining of dissimilar materials is becoming increasingly prevalent to combine differing material properties to enhance product design flexibility and performance. In this study, pure copper was built on 316L stainless steel (SS316L) by laser-based additive manufacturing technology in which copper was deposited layer-by-layer on SS316L with Inconel 718 intermediate layers.

The goal is to fabricate multi-metallic structures with improved thermal conductivity. The direct joining of Cu on SS316L would result in porosities at the interface and the consequent poor mechanical properties, which could be addressed by Inconel 718 intermediate layers. The microstructure, chemical composition, tensile properties, and micro-hardness were characterized in the dissimilar materials using scanning electron microscopy, energy dispersive spectroscopy, tensile test with digital image correlation technique, and hardness tester. Results confirm excellent bonding when Inconel 718 intermediate layers are introduced.Keywords: Additive Manufacturing; Functionally Graded Materials; Cu; Stainless Steel; Dissimilar materialsIntroductionMany metallic components in modern systems are frequently subjected to a complex and diverse working environment.

For example, a jet engine has a large number of parts that are subjected to severe working conditions including high temperature, high pressure, corrosion, strong centrifugal load, etc. [1]. Meanwhile, the working condition is location-dependent, which indicates different parts can experience varied conditions. Therefore, a single part that has location- dependent composition and thus properties is highly desired in future advanced systems to offer distinct properties within the single part, which can better fulfill the design requirements [2]. Such a desire promotes the design and fabrication of functionally graduated materials (FGM). FGM is a novel material that features different compositions in different positions, resulting in corresponding changes in the properties in terms of mechanical, thermal, electrical, chemical, etc.

FGM is a novel material that features different compositions in different positions, resulting in corresponding changes in the properties in terms of mechanical, thermal, electrical, chemical, etc. FGM has the benefits to combine the strength of each material and thus provide a unique solution for the design of components that need to work in a complex environment.With the emerge of many advanced manufacturing techniques, the fabrication of FGM becomes possible. Among these advanced manufacturing techniques, additive manufacturing (AM) is a preferred choice for making FGM due to its fabrication nature. According to the American Society for Testing and Materials (ASTM) F2792-12a, AM can be classified into seven categories based on the manufacturing principle [2]. Many of these technologies are particularly being1011investigated, such as powder bed fusion (PBF), binder jetting, and directed energy deposition (DED).

Many of these technologies are particularly being1011investigated, such as powder bed fusion (PBF), binder jetting, and directed energy deposition (DED). For the fabrication of FGM structures, DED is particularly suitable. In the DED process, materials in the form of powder or wire are deposited on a substrate according to a designed tool path. The shape of an object is obtained in a layer-by-layer fashion, which is opposite to the traditional machine process where materials are cut from a block. Since materials are added layer upon layer, the material composition can be changed as a function of layer by easily changing the feedstock material. The material composition can even be changed as a function of a spot within a single layer. Such flexibility makes DED widely used to fabricate FGM structures. Based on engineering needs, a wide type of FGM has been designed and made by AM. For instance, Li et al.

Such flexibility makes DED widely used to fabricate FGM structures. Based on engineering needs, a wide type of FGM has been designed and made by AM. For instance, Li et al. designed and fabricated Ti6Al4V-SS316L (stainless steel 316L) FGM to combine the high strength-to-weight ratio and high-temperature strength/toughness of Ti6Al4V with the corrosive resistance and low cost of SS316L [3]. A special grading route Ti6Al4V -> V -> Cr -> Fe -> SS316L was designed to eliminate the formation of detrimental brittle phases. Inconel 625 was joined with SS316L to take advantage of the high strength at elevated temperatures which Inconel 625 can provide and good machinability and lower price which SS316L can provide [4].

Inconel 625 was joined with SS316L to take advantage of the high strength at elevated temperatures which Inconel 625 can provide and good machinability and lower price which SS316L can provide [4]. Many other types of FGM have been reported including Inconel 718 – GPCop 84 copper [5], Ti6Al6V- Invar [6], SS304L-Inconel 625 [6], and H13 tool steel – copper [7].Joining copper with stainless steel (SS) to create copper-SS FGM is very attractive due to the wide application of this structure. Copper is known for its very high thermal and electrical conductivity. Therefore, joining copper with SS can simultaneously use the high thermal conductivity of copper and the high strength and corrosion resistance of SS. Such a hybrid structure can be used in places where efficient cooling is desired, such as in power generation industries, nuclear industries, heat exchangers, die casting parts, etc [8].

Such a hybrid structure can be used in places where efficient cooling is desired, such as in power generation industries, nuclear industries, heat exchangers, die casting parts, etc [8]. Traditionally, copper was joined with SS by the welding process, such as explosive welding [9], laser welding [10], and diffusion welding [11]. It should be noted that since copper and stainless steel have significant dissimilarities in physical, chemical, and thermo-mechanical properties, joining copper with stainless steel is a challenging task. It is reported that Fe and Cu have very limited solid solubility [12]. In addition, high residual stress can be built up in the welding interface, which can cause cracking [13]. Previous studies have reported the failure in joining copper with stainless steel due to the aforementioned drawbacks [14][15][16][17].This study aims to investigate the feasibility of fabricating copper-SS FGM structures using additive manufacturing.

Due to the crack-prone phenomenon in the direct joining of copper on SS, intermediate layers made of Inconel 718 (IN718) was added between copper and SS. IN718 has a high content of element Ni and since Ni can form a solid solution with Cu, it is expected that the final FGM is defect-free. In this study, SS316L was deposited first, followed by the deposition of IN718 on SS316L, and finally followed by the deposition of copper on IN718. The resulting structure was characterized in terms of microstructure, elemental distribution, tensile properties, and hardness distribution. The outcome of this study provides an innovative solution to join copper with stainless steel using additive manufacturing techniques to enhance the overall performance of the FGM structure.1012Feedstock Materials and Experimental SetupSS316L, IN718, and Cu powder materials were provided as the feedstock. The SS316L and IN718 powders were procured from Carpenter Technology Corporation.

The SS316L and IN718 powders were procured from Carpenter Technology Corporation. The Cu powder was supplied by Royal Metal Powders Inc. All three types of powders were fabricated by the gas optimization process. Figures 1a, 1b, and 1c depict the scanning electron microscopy (SEM) images of SS316L, IN718, and Cu powder, respectively. As can be seen, the micro-topographies of these three types of powders are spherical while some tiny satellite particles are scarcely found and attached to major large particles. ImageJ software was used to characterize the particle size and distribution of the powders. As can be seen in Figures 1d, 1e, and 1f, these three types of powders are featured with normal distribution with an average particle diameter of 116 µm for SS316L, 79 µm for IN718, and 110 µm for Cu powder. The chemical composition of SS316L, IN718, and Cu powder feedstock was provided by the powder manufacturers and the composition is displayed in Table 1.

The energy distribution of the circular laser beam follows the Gaussian profile. The relative movement between the laser head and deposition platform is realized by a 3-axis stage. An electrostatic powder feeder (X2 powder feed system from Powder Motion Labs, USA) was used in this study for material feeding. Argon is utilized as the carrier gas to deliver powder particles from the powder feeder to the molten pool. Argon is also used as shielding gas in the deposition enclosure to prevent materials from oxidation. All DED components were controlled by a customized computer numerical control (CNC) software developed in Linux. The DED deposition principle is schematically illustrated in Figure 2b. The laser beam is in the vertical direction and the powder feed nozzle is 20° tilted from the laser beam. The distance from the end of the powder nozzle to the top surface of the substrate is 15 mm. During deposition, the laser beam creates a molten pool on a substrate.

The distance from the end of the powder nozzle to the top surface of the substrate is 15 mm. During deposition, the laser beam creates a molten pool on a substrate. Meanwhile, powders are transported into the molten pool, melted, and then solidified to form a laser. This process continues so that the designed 3D shape can be formed.Figure 2. Schematic of (a) DED equipment and (b) deposition process.The multi-metallic structure as presented in Figure 3a was fabricated in this investigation. First, an SS316L thin-wall structure with a length of 36 mm and a height of 8 mm was deposited on an SS304L substrate. The layer thickness is set as a constant value of 0.5 mm. The deposition1014Cu8.96 1085 400 17tool path is a single-track multi-layer pattern as inserted in Figure 3a. Materials were deposited in both forward and backward directions with no interlayer dwell. This tool path was also applied to the deposition of IN718 and Cu as discussed below.

Materials were deposited in both forward and backward directions with no interlayer dwell. This tool path was also applied to the deposition of IN718 and Cu as discussed below. Once the SS316L was deposited, the as- deposited SS316L and substrate were cooled to room temperature. Next, IN718 was deposited on the top of SS316L. The part was cooled down to room temperature again before the deposition of Cu on the top of IN718. The detailed processing parameters are given in Table 3.Table 3. DED processing parameters for fabricating the SS-IN718-Cu FGMParameterLaser power (W) Laser beam size (mm) Scan speed (mm/min) Layer thickness (mm) Number of layers Deposit length (mm) Deposit height (mm)SS316L 600 2 210 0.5 16 36 8SS-IN718-Cu FGMIN718 500 2 210 0.4 6 36 2.4Cu 550 2 210 0.3 16 36 4.8Figure 3.

(a) Schematic of SS-IN718-Cu FGM; (b) Sectioned vertical cross-section; (c) Dimensions of tensile test specimens (all dimensions are in mm).A vertical cross-section was sectioned from the as-fabricated structure as shown in Figure 3b. The cut piece was mounted on conductive Bakelite. After that, the sample was ground using silicon carbide abrasive papers from 120 grit to 1200 grit in sequence, polished using 9 µm, 6 µm, 3 µm, and 1 µm diamond suspension, and finally polished using 0.02 µm silica suspension. The1015sample was ultrasonically cleaned immediately after polishing. Since the microstructure of different materials is of interest, different etchants were selected. In order to show the microstructure of IN718 and SS316L, the sample was firstly electrolytically etched with 10% oxalic acid under 2V for 30 seconds.

In order to show the microstructure of IN718 and SS316L, the sample was firstly electrolytically etched with 10% oxalic acid under 2V for 30 seconds. After that, the sample was etched using a solution of 70% ethanol, 25% hydrochloric acid, and 5% ferric chloride by immersion for 15 seconds to reveal the microstructure of the Cu. The prepared sample was analyzed using scanning electron microscopy (SEM, FEI Helios NanoLab 600 Dual Beam) coupled with an Oxford Energy Dispersive X-ray Spectroscopy (EDS). EDS mapping was carried out on the SS-IN718-Cu FGM to study elemental diffusion across the interface of dissimilar materials.In order to reveal the mechanical properties of the SS-IN718-Cu FGM, tensile specimens were prepared from the as-fabricated sample as illustrated in Figure 3a. Tensile specimens were cut in the vertical direction. The dimensions of tensile specimens are shown in Figure 3c. The tensile gauge section consists of three materials (SS, IN718, and Cu).

Tensile specimens were cut in the vertical direction. The dimensions of tensile specimens are shown in Figure 3c. The tensile gauge section consists of three materials (SS, IN718, and Cu). Therefore, tensile testing of these samples can reveal the bonding strength between dissimilar materials. The tensile specimens were tested using an Instron universal testing machine with a crosshead speed of 0.15 mm/min. In-situ digital image correlation (DIC) technique was used to show the strain maps during the tensile test. The longitudinal surface of the fractured sample was analyzed using SEM to study the region of fracture. Vickers microhardness test was also performed on the SS-IN718-Cu FGM to investigate hardness profile across different layers. Vickers microhardness measurements were carried out along the vertical building direction using a Struers Druamin-5 hardness tester with a load of 4.9 N and a hold time of 10 seconds.

Vickers microhardness measurements were carried out along the vertical building direction using a Struers Druamin-5 hardness tester with a load of 4.9 N and a hold time of 10 seconds. Three measurements were conducted at the same height, and the average value and standard deviation were reported.Results and DiscussionMicrostructureThe microstructure of the SS-IN718-Cu FGM at different locations is shown in Figure 4. As can be seen from Figure 4a, the SS316L layers were dominated by columnar dendrites. This columnar structure was widely reported previously in the as-deposited SS316L materials both manufactured by directed energy deposition [18] and selective laser melting process [19]. In SS316L deposits close to the substrate, columnar dendrites usually grow in the vertical direction. This is because the temperature gradient is significant in the vertical direction, causing the dendrites to elongate in this direction.

This is because the temperature gradient is significant in the vertical direction, causing the dendrites to elongate in this direction. With the increase of layers, SS316L tends to transform to a mixture of columnar dendrites and equiaxed structures. This variation of microstructure in different layers of additively manufactured SS316L can be explained by the solidification theory [20]. In the solidification process, the solidification velocity (V) and temperature gradient (G) are important parameters to guide microstructure formation. With the reduction of G/V, the microstructure can change from planar to dendrite and finally to an equiaxed structure. At the beginning of the deposition process (initial layer), the temperature of the substrate is relatively low. This can result in a very high temperature gradient between the as-deposited layers and substrate. It is reported that this temperature gradient G can reach up to 8 × 105 K/m in the deposition of SS316L [21].

It is reported that this temperature gradient G can reach up to 8 × 105 K/m in the deposition of SS316L [21]. Therefore, the G/V ratio can be as high as 900 – 6000 Ks/mm2 as reported in [22]. This extremely high G/V ratio caused the formation of columnar dendrites in the vertical direction in the lower SS316L layers. With the increase of layers, the temperature gradient between the1016newly deposited layer and the layer underneath is reduced. Since a relatively low G/V was achieved in the upper layers, the upper layers are featured with a mixture of columnar and equiaxed structures.Figure 4. Microstructure of (a) SS316L close to the substrate, (b) IN718-SS316L transition zone, (c) middle layer of IN718, and (d) Cu-IN718 transition zone.Figure 4b shows the microstructure at the SS316L and IN718 transition zone. In terms of defects, no defects such as delamination or pores were identified in the interface.

In terms of defects, no defects such as delamination or pores were identified in the interface. The microstructure at the transition zone shows a good bonding between SS316L and IN718. SS316L and IN718 are both fcc (face-centered cubic) austenite structures. Based on the Fe-Ni phase diagram as seen in Figure 5a, a large window with no detrimental intermetallic can be achieved. It is expected that a solid bonding can be obtained between SS316L and IN718, which can be validated by the microstructure analysis in Figure 4b and mechanical testing discussed in the following of this paper. In terms of microstructure, the IN718 featured with different structure when compared with the SS316L below the interface, as can be seen in Figure 4b. Such difference1017in microstructure can also be explained by the solidification mechanism discussed above.

Such difference1017in microstructure can also be explained by the solidification mechanism discussed above. As mentioned in the experimental procedure section, the as-deposited SS16L and substrate were cooled down to room temperature before the deposition of IN718. Therefore, in the deposition of the first layer of IN718, the G/V is extremely high, causing the formation of columnar structure in the direction where the temperature gradient is the highest. The microstructure of IN718 in the middle layer also reveals dominated columnar dendrites (Figure 4c). Since the IN718 was deposited as intermediate layers with a total layer number of 6, it is reasonable to expect that the G/V is still high in the middle and upper IN718 layers, resulting in the columnar dendritic structure.Figure 5. Phase diagrams for (a) Fe-Ni and (b) Cu-Ni system.Figure 4d shows the microstructure at the IN718-Cu transition zone. The micrograph confirms that the IN718-Cu interface is defect-free.

Phase diagrams for (a) Fe-Ni and (b) Cu-Ni system.Figure 4d shows the microstructure at the IN718-Cu transition zone. The micrograph confirms that the IN718-Cu interface is defect-free. A good interfacial bonding strength can be expected (validated by tensile testing below). Since the Cu was >99.99 % pure, no obvious structure was observed in the bulk Cu deposits. However, the deposits close to the interface featured unevenly distributed composition and structure as can be seen in Figure 4d. This has resulted from the diffusion of composition from the IN718 layers to the layers above the interface. Such phenomenon is discussed below supported by EDS data.Figure 6 shows the EDS maps over the IN718-Cu transition zone. It reveals that above the interface, Cu is the dominant content which is expected. However, areas, where IN718 was rich, can also be found. This shows that the IN718 composition penetrated the Cu layers.

However, areas, where IN718 was rich, can also be found. This shows that the IN718 composition penetrated the Cu layers. A gradual transition of composition across the interface was observed due to large elemental diffusion. This chemical dilution is common in additively manufactured multi-material structures. In the additive manufacturing process, when the lower material is melted (IN718 in this study), the incoming new material (Cu in this study) is delivered into the melt pool and mixed with the existing material. When the material is solidified, a composition of the lower material can be found in the newly deposited layers. In the additive manufacturing process, the Marangoni convection happens in the molten pool. The thermal Marangoni convection is positively proportional to surface tension and temperature gradient. In the deposition of Cu on IN718, the center of the molten pool has a higher temperature than the edge.

In the deposition of Cu on IN718, the center of the molten pool has a higher temperature than the edge. This temperature difference causes a surface tension gradient toward the edge of the molten pool. Therefore, the lower material is pulled into the upper location. In this study, the pure Cu has a very high thermal conductivity, which can cause intensified dynamics of the Marangoni convection in the molten pool. Such intensified Marangoni convection caused the1018large dilution of IN718 into the Cu layers. It should be mentioned that large interfacial dilution is helpful for residual stress mitigation since the discrepancy between two distinct materials was wakened [23].Figure 6. Elemental mapping over Cu-IN718 transition zone.Figure 7. Elemental mapping over IN718-SS316L transition zone.1019Figure 7 shows the EDS maps across the SS-IN718 transition zone. An interface where the material starts to change can be seen.

Elemental mapping over IN718-SS316L transition zone.1019Figure 7 shows the EDS maps across the SS-IN718 transition zone. An interface where the material starts to change can be seen. The material below the interface is rich in Fe, aligned with the fact that more Fe in SS316L. The Cr composition is similar above and below the interface since very close of Cr composition in SS316L and IN718, as can be seen in Table 1.Tensile PropertiesAs illustrated in Figure 3a, each tensile specimen has the as-deposited Cu, IN718, and SS316L. Tensile testing of the specimens can provide an understanding of the bonding strength between dissimilar materials. Figure 8 presents a representative tensile stress-strain curve. Table 4 summarizes the 0.2% offset yield strength (YS) and ultimate tensile strength (UTS) of this study compared with existing data. The result shows the average YS and UTS are 155 ± 7 MPa and 257 ± 25 MPa, respectively.

The result shows the average YS and UTS are 155 ± 7 MPa and 257 ± 25 MPa, respectively. Figure 9 shows the in-situ DIC strain field at different stages in the testing of the specimen. DIC result shows that deformation has mostly happened at the Cu section, which is expected since the strength of Cu is the lowest among the Cu, IN718, and SS316L. The low tensile strength was not able to cause significant deformation over the IN718 and SS316L regions. All the tensile specimens fractured at the Cu region, demonstrating the bonding strength of the multi-material system is higher than the UTS of the Cu. Table 4 lists the previously reported tensile properties of pure or almost pure Cu made of additive manufacturing or traditional process techniques. The data shows the tensile strength of the Cu made in this study is comparable to the literature data.Figure 8. Stress-strain curve of Cu-IN718-SS316L FGM.1020Table 5.

In-situ DIC strain field during tensile testing of Cu-IN718-SS316L specimens (percent indicates strain percentage).Vickers HardnessThe microhardness was measured from the substrate to the as-deposited Cu along the vertical building direction. Figure 10 shows the hardness profile. The measurement reveals that the as-fabricated SS316L featured a hardness of 173 ± 6 HV. This is a typical value compared with SS316L made by additive manufacturing [30]. When reaching the IN718 region, the hardness shows a gradual increase. It can be seen that since elemental diffusion occurred in the IN718-SS transition zone, the hardness shows a gradual increase with the distance approaching the IN718 region. The as-deposited IN718 has a maximum hardness of 253 HV at approximately the middle of the IN718 layers. This is because of the composition dilution at the SS-IN718 and IN718-Cu transition zones, which weakened the hardness of IN718.

This is because of the composition dilution at the SS-IN718 and IN718-Cu transition zones, which weakened the hardness of IN718. Extending to the Cu region, a gradual1021drop of hardness is seen. Since Ni can form a solid solution with Cu (Figure 5b), the diffusion of IN718 into the Cu layers is significant, causing the gradual decrease of hardness. The Cu upper layers have a roughly stabilized hardness of 70 HV.Figure 10. Hardness distribution along the building direction of the SS-IN718-Cu FGM.ConclusionThis study investigated the fabrication of functionally graded structures by joining pure copper with stainless steel via Inconel 718 intermediate layers by an additive manufacturing process. The ultimate goal of this study is to combine the excellent thermal conductivity of copper with the high strength and corrosion resistance of stainless steel for applications where efficient cooling is required.

In the present work, the SS316L-IN718-Cu FGM structure has been successfully fabricated. The final FGM structure was evaluated in microstructure, tensile properties, and microhardness distribution. The primary findings are:(1) Defect-free FGM structure has been achieved by inserting IN718 interlayers between SS316L and pure copper. IN718 intermediate layers successfully eliminated the cracking issue in the direct joining of SS316L with copper.(2) A transition of microstructure was observed along the building direction. This change of microstructure can be explained by the solidification theory guided by the change of G/V. Columnar grains dominated the lower SS316L layers while the upper SS316L layers have a mixture of columnar and equiaxed structure. The IN718 layers are dominated by columnar structure.(3) Elemental diffusion happens across both the SS-IN718 and IN718-Cu interfaces. This extended elemental diffusion is due to the Marangoni convection in the molten pool.

This extended elemental diffusion is due to the Marangoni convection in the molten pool. (4) The YS and UTS of the SS316L-IN718-Cu structure are 155 ± 7 MPa and 257 ± 25 MPa, respectively. The FGM structure fractured at the Cu region while the interfaces are survived, indicating the bonding strength is higher than the UTS of the FGM.1022(5) A change of hardness is observed along the building direction due to material differences. Hardness varies smoothly across the interfaces due to large elemental dilution.ConclusionThis project was supported by National Science Foundation Grants CMMI-1547042 and CMMI 1625736, Intelligent Systems Center, Center for Aerospace Manufacturing Technologies, and Material Research Center at Missouri S&T. Their financial support is greatly appreciated.References[1] X. Zhang, W. Li, F. Liou, Damage detection and reconstruction algorithm in repairingcompressor blade by direct metal deposition, Int. J. Adv. Manuf. Technol. 95 (2018) 2393–2404.

ditive Manufacturing of Directionally Heat Conductive ObjectsKulkarni P.M*, Karunakaran K.P*, Asim Tewari*, Fisseha Legesse*, Dhirendra Rana*, Alain Bernard**Rapid Manufacturing Laboratory, Department of Mechanical Engineering, Indian Institute of Technology Bombay, Powai, Mumbai-400076, INDIA **Virtual engineering for industrial engineering, IRCCyN, Ecole Centrale de Nantes, FranceAbstractPurpose: Directionally Heat Conductive (DHC) objects are also known as thermal cables. DHC objects used in cooling solid state devices as large heat sinks. This paper presents Additive manufacturing (AM) as a new method for realization of DHC objects. AM can produce DHC objects accurately, rapidly and economically. Design/Methodology/Approach: It is very hard to realize DHC objects through conventional manufacturing. The authors describe the realization of metallic DHC objects using their AM process called Hybrid Layered Manufacturing (HLM).

The authors describe the realization of metallic DHC objects using their AM process called Hybrid Layered Manufacturing (HLM). HLM, which uses GMAW deposition, can, in principle, produce DHC out of any metal for which welding wire is available. . Differential conductivity in HLM is achieved by appropriately distributing the air gaps among the beads. The air gap distribution depends on the layer thickness and step over increment. Findings: The authors have demonstrated the ability of HLM to produce DHC objects for Al alloy 4043. The thermal conductivity achieved on two orthogonal directions was 100.496 and 129.740 W m-1 K-1 as against 163 W m-1 K-1 for solid metal. These are significant variations. Research Limitations/Implications: HLM realize near net shape of DHC objects, further post processing namely machining is required.

These are significant variations. Research Limitations/Implications: HLM realize near net shape of DHC objects, further post processing namely machining is required. Originality Value: AM Specifically HLM for realization of DHC objects Keywords: Directional heat conductive (DHC), Additive Manufacturing (AM), Hybrid Layered Manufacturing (HLM).1. IntroductionThermal conductivity of a material has great influence on its performance in heat transfer analysis, a material is normally assumed to be isotropic. Therefore, it is a common practice to evaluate the thermal conductivity at the average temperature and treat it as a constant in calculations [1]. This assumption is realistic for most materials, but some materials, such as laminated composites, are anisotropic. This anisotropy can be engineered to have directional conductivity. Such objects are known as Directionally Heat Conductive (DHC) objects.

This anisotropy can be engineered to have directional conductivity. Such objects are known as Directionally Heat Conductive (DHC) objects. There are several applications of DHC in different fields.The characteristics of a DHC object depend on the material composition and porosity or void along different directions. DHC objects find major applications as heat sinks in electronic packaging, especially for high voltage chips, packaging base plates, carrier plates of electronic box closures [2, 3]. In these applications, heat dissipation is important because modern integrated circuits and other powerful electronic parts generate a lot of heat.H. Amiri et al. have reported the numerical solution for the steady state combined conductive-radiative heat transfer in an anisotropic participating medium within the irregular geometries [4].

Amiri et al. have reported the numerical solution for the steady state combined conductive-radiative heat transfer in an anisotropic participating medium within the irregular geometries [4]. Some of the engineering applications of such combined properties were mentioned in areas such as thermal transport in nano-materials, heat transfer through the semi- transparent, porous materials, multi-layered insulations, glass fabrication etc. [5]. Similarly DHC objects can be used as effective heat conductors in large, high speed electrical machines as housing jackets to extract the heat generated inside the stator. It can be also used for insulation in industrial furnaces to retain the heat [6]. Aerospace engines uses honeycomb structures which provides directional thermal conductivity as well as weight reduction.Temperature rise in high speed electrical machine causes mechanical failures and irreversible demagnetization of permanent magnets.

Initially temperature in high speed electrical machines was maintained by using forced air-cooling with external cooling fins often by a ventilator mounted on the shaft. At higher speeds, this may be very noisy and take up a fair deal of torque, so an independent cooling is required. This independent cooling is done by DHC housings.Figure 1 shows the current methods of manufacturing DHC objects like bonding, sheet metal forming (folding and co-corrugation process), open and closed die forging, liquid forging, skiving, Powder Metallurgy (PM), Metal Matrix Composite (MMC) manufacture [7]. DHC objects manufactured by various process is shown in figure 1.Figure 1 Various Manufacturing Methods of DHC Objects[7]Honeycomb structure is manufactured by sheet metal forming method named as co- corrugation process. The unidirectional thermally conductive fibers are oriented in a direction parallel to the width of the formed sheet.

The unidirectional thermally conductive fibers are oriented in a direction parallel to the width of the formed sheet. There are various configurations of honeycomb structures, out of which hexagonal cells gives better directional thermal conductivity. Titanium honeycomb structure used for exhaust nozzles in aircraft is shown in figure 3(a) and this type of structure is manufactured by corrugation process shown in figure 3(b). This honeycomb structure finds major applications in military and commercial aircraft, automotive bodies, engine components, recreational equipment, and marine craft [8, 9].Figure 2 DHC objects Manufactured by various Process [7](a) Metallic Honeycomb(b) co-corrugation process Figure 3 Manufacturing of DHC Objects by Sheet Metal Forming Process[8]MMC used to manufacture DHC vibration isolator imparting shock and vibration isolation. MMC can be manufactured by powder metallurgy and foil diffusion, both of them are solid based methods.

MMC can be manufactured by powder metallurgy and foil diffusion, both of them are solid based methods. Electroforming, stir casting, squeeze casting, spray deposition and reactive processing are the liquid based MMC manufacturing processes [10].Heat sinks are manufactured by forging instead of casting, machining and extrusion. Forging is the most effective method to form complex shapes and also offers unique thermal advantages.The part is formed under high pressure which controls the grain structure and results in improved thermal performance. The cold forging process results in a heat sink that yields a 14% improvement over the extruded and a 62% improvement over the die cast heat sink [11].The above methods to produce DHC objects are hard and expensive as they are dominated by subtractive manufacturing (AM), manufacturing of DHC objects has become elegant and affordable. AM is inherently anisotropy.

AM is inherently anisotropy. When this inherent anisotropy is engineered to meet the functional needs of a component, it is known as Functionally Gradient Materials (FGM). DHC is a special case of FGM.[12]. Withthe advent of Additive ManufacturingAdditive Manufacturing (AM) is the term given to a group of technologies that are capable of creating physical objects from computer aided design (CAD) files by incrementally adding material such that the objects “grow” from nothing to completion. Most of today’s AM technologies start with a CAD solid model which is sliced into thin layers. Each layer comprises a 2D cross section profile of the part which is then made layer by layer.AM involves a number of steps that move from the virtual CAD description to the physical resultant part. Different products will involve AM in different ways and to different degrees.

Different products will involve AM in different ways and to different degrees. Small, relatively simple products may only make use of AM for visualization models, while larger, more complex products with greater engineering content may involve AM during numerous stages and iterations throughout the development process. Furthermore, early stages of the product development process may only require rough parts, with AM being used because of the speed at which they can be fabricated. At later stages of the process, parts may require careful cleaning and post-processing (including sanding, surface preparation and painting) before they are used, with AM being useful here because of the complexity of form that can be created without having to consider tooling [7].The automated approach to creating objects from CAD files in the additive manner that AM allows is beginning to have a profound effect on how objects are made.

Almost all manufactured objects that we see today are made with some form of tooling at some point – parts may be machined using a cutting tool or parts may be formed from a molding tool. The use of tooling imposes numerous restrictions such as design limitations, restrictive costs to prepare for manufacture and the need to create products at a single location prior to being shipped globally. AM technologies invariably require no tooling to create parts and this profoundly affects how products can be designed, tailored, made, distributed etc. [12].Organization of PaperIn this paper authors give the introduction of HLM process, involvement of HLM technology in realization of DHC objects, characterization of DHC objects by using the heat conductivity testing. Finally authors give the conclusion in which they show unidirectional behavior of DHC objects.2.

Finally authors give the conclusion in which they show unidirectional behavior of DHC objects.2. Hybrid Layer Manufacturing ProcessHybrid Layered Manufacturing (HLM) is a Rapid Manufacturing process which includes layer by layer deposition using Gas Metal Arc Welding (GMAW). This deposition makes near-net shape of the object to be built which is then finish machined subsequently. Face-milling is doneafter depositing each layer to remove oxidized surface and scallops. It also ensures z-accuracy of each layer. Integration of HLM with 5 axis and 3 axis CNC machines is shown in figure 4.Arc weld deposition is faster and more economical by an order of magnitude than laser and electron beam processes (50-130g/min as against 2-10g/min). The surface quality is in the decreasing order for laser, electron beam and arc.

The surface quality is in the decreasing order for laser, electron beam and arc. Even the best surface obtained through AM is not adequate enough for direct use for very serious applications except in some medical applications where rough surface is favourable for tissue growth. If finish machining is necessary, then these marginal quality differences do not matter.Subtractive processes can produce good quality parts but are slow; although the material removal by itself is fast, human efforts required for cutter path generation is the bottleneck. On the other hand, additive processes are fast but produce poor quality parts. The additive process focuses on speed while ensuring the desired material integrity. The resulting object is only near- net as no attention to the geometric quality is paid at the time of building it in layers. In HLM, the near-net shape is obtained with total automation.

The resulting object is only near- net as no attention to the geometric quality is paid at the time of building it in layers. In HLM, the near-net shape is obtained with total automation. However, the subsequent two stages of stress relieving/ heat treatment and finish machining have fair amount of human intervention. In other words, HLM does not strive to achieve total automation but aims at optimal/ economical level of automation.Welding Gun(b) Close view of 5 Axis CNC HLM unit(a) 5 Axis CNC machineFace Milling CutterWelding Gun(c) 3-Axis CNC machine(d) Close view of 3-Axis HLM unitFigure 4 Integration of HLM with conventional CNC machines2.1 Methodology of HLMThe stages involved in building the metallic object using HLM is shown in figure 5.a. Geometric processingb. Building the near-net shapec. Stress relievingd.

Geometric processingb. Building the near-net shapec. Stress relievingd. Finish machiningGeometric Processing CAD Model STL Format Slicing CNC ToolpathBuilding Near - Net Shape MIG Weld Deposition Surafce MillingStress RelievingFinish MachiningFigure 5 Process FlowchartA case study of impeller which was manufactured using HLM process is shown below infigure 6.2.2 ApplicationThe various geometric capabilities of HLM, in its three kinematic modes of 3-axis, positional 5- axis and continuous 5-axis, are shown in figure 7. In 3 axis kinematics mode, we can manufacture parts by means of various approaches such as planar deposition, planar deposition with small undercuts, planar deposition after building the undercut, non-planar (conformal) deposition.

HLM can be used to manufacture Functionally Gradient Materials (FGM) by using multiple welding torches at with different filler materials.2.3 Advantages of HLMi.Seamless integration of material addition and subtraction on the same platformii. Totally automatic manufacture of the near-net shapeiii. Use of economical, faster and safer GMAW as against laser and electron beamiv. Retrofittable to any existing CNC machine, independent of make and agev. Easy toggling between HLM mode and regular CNC machining mode(a) CAD Model(b) CAD Model after slicing(c) Near net shape(d) Finished partFigure 6 Case Study Shows HLM ProcessFigure 7 Capabilities of HLM in Different Kinematic Modes [15]3. Realization of DHC Using HLMA Cold Metal Transfer (CMT) 2700 welding machine and a 3-axis CNC machine were used for deposition and subtraction respectively.

Realization of DHC Using HLMA Cold Metal Transfer (CMT) 2700 welding machine and a 3-axis CNC machine were used for deposition and subtraction respectively. Three axis Cartesian coordinate system was adopted to identify the direction of deposition and directionally heat conductivity analysis. The X-axis (torch direction) is taken in the long travel of the weld torch, Z-axis is taken perpendicular to the surface of the substrate (weld direction) and Y-axis is taken in the stepover increments direction. Figure 8 illustrates the coordinate system used in deposition and characterization. The heat conduction was measured in the directions parallel and perpendicular to the air gaps in X and Y directions respectively.The part was built by adopting zig-zag deposition. The deposition in the longitudinal direction was built on X-axis and the pitch was provided in a perpendicular direction. The pitch size was given by considering the air gaps between consecutive beads.

The pitch size was given by considering the air gaps between consecutive beads. Complete fusion between weld beads was prevented by providing a step over which is equal to half of the bead width plus the required air gap. The CAD model of the sample, with air gaps between each strip of layers, is shown below in figure 8 (d).(a) Front View of Specimen(b) Top View of specimenAir Gaps(c) Coordinate system for deposition and characterization(d)CAD model of a DHC sampleFigure 8 Design of DHC SampleAs shown in Table 1, during deposition different process parameters were taken into account. Wire speed which affects the bead width and thickness, and which in turn is affected by the rate of heat input, was controlled to minimize distortion of beads due to overheating during layer deposition. The torch speed which is another important welding parameter was controlled numerically to get a high precision in the weld bead size and to minimize heat distribution over a wider area [13].

The torch speed which is another important welding parameter was controlled numerically to get a high precision in the weld bead size and to minimize heat distribution over a wider area [13]. If the torch speed is high, the bead width is thin and vice-versa. Therefore thelowest wire speed and the highest torch speed within their stable operating ranges are chosen to preserve the required air gap. The other and the most important parameter which was critical to our specific experiment is the stepover increment. The size of the stepover determines the size of the air gaps. If the stepover is less than half of the width of the weld bead, complete fusion between adjacent weld beads will occur. The size of stepover was determined based on the required size of air gaps.Variation of current with layer thickness is shown in figure 9.

The size of stepover was determined based on the required size of air gaps.Variation of current with layer thickness is shown in figure 9. It shows that as currentincreases layer thickness also increases.Figure 9 Variation of Layer thickness with currentThe process parameters used during weld deposition are shown in table 1. The parametersare given based on the requirements of the final object.Table 1 Welding Parameters for Al-Si Layer thickness (mm) Torch speed(mm/min) Wire feed rate(m/min) Welding wire diameter (mm) Gas flow rate(l/min) Shielding gas Pitch Air gaps1 850 3.4 1.2 10-12 Ar Bead width+ 0.25mm 0.25mm4. Characterization of DHCCharacterization of DHC includes the thermocouple calibration which is used to plot a graph between temperature and voltage using which an equation is established.

Characterization of DHCCharacterization of DHC includes the thermocouple calibration which is used to plot a graph between temperature and voltage using which an equation is established. Furthermore a heat conduction test is performed to show the unidirectional heat flow.4.1 Thermocouple CalibrationThe necessity of the calibration is to establish an equation of curve that fits to the temperature versus voltage graph. This equation is applied to find out the unidirectional temperature drop orrise in the sample along the specified directions. Hence a calibration setup was established to calibrate a thermocouple based on a manual by ASTM [14] using fixed reference temperature technique. If properly used, it is a convenient and inexpensive way to achieve the ice point, it can be reproduced with ease and with an exceptional accuracy of 0.0001oC [14]. Ice bath reference junction setup was established for this purpose as shown in figure 10.

Ice bath reference junction setup was established for this purpose as shown in figure 10. The ice bath is made up of a mixture of melting shaved ice and water. This is one of the most common reference junction techniques used in calibration. To avoid the errors due to melting of the ice at the bottom of the bath, consumed excess water was removed periodically and more ice was added, so that the ice level is maintained safely below the reference junctions.The water level in the bath was checked and the apparatus was connected to a power supply. The temperature in the bath was initially set to 30°C to start heating. In normal mode, the temperature shown on the digital display is the actual temperature of the bath (TB) against which the thermocouple sensor is to be calibrated. The ends of the thermocouple are connected to the digital multimeter and which was set to read voltage in mV. The other end of the thermocouple was dipped in ice bath.

The ends of the thermocouple are connected to the digital multimeter and which was set to read voltage in mV. The other end of the thermocouple was dipped in ice bath. A stable temperature source was kept to maintain a constant temperature. The multimeter reading in millivolts was noted down for each 5°C increment in temperature from 30°C to 80°C. The most commonly used reference temperature is 0oC [13].The multimeter reading in millivolt corresponding to the difference in temperature between the surroundings (ice) and the bath was recorded. The measured bath temperatures values (TB) were plotted on Y-axis against the corresponding reading of emf (mV) values on X-axis, as shown in Fig.13. The slope and intercept of the curve-fitted were obtained from the equation correlated to the graph.

The slope and intercept of the curve-fitted were obtained from the equation correlated to the graph. The linear correlation coefficient (R), which measures the strength and the direction of a linear relationship between two variables, was calculated to observe the strength of association between the millivolt reading and the temperature during calibration.R=𝑛(∑𝑇∗𝑉)−(∑ 𝑇)∗(∑𝑉) √𝑛(∑ 𝑇2)−(∑ 𝑇)2 √𝑛(∑ 𝑉2)−(∑ 𝑉)2Where n is number of pairs of data T is temperature in 0C V is voltage in mVFigure 10 Thermocouple calibration setupThe thermocouple voltage at various known temperatures was measured. This will develop acorrelation between thermocouple voltage and thermocouple temperature. This correlation is represented by the graph in figure 11.

This will develop acorrelation between thermocouple voltage and thermocouple temperature. This correlation is represented by the graph in figure 11. As it can be seen from the graph that the variation in temperature corresponds to the variation in voltage reading.Figure 11 Correlation between temperature and voltage readingsThe end result of regression analysis predicts the value of an output variable (temperature) based upon the values of input variables (mV reading). So based on the result of the regression which is 99.6% the total variation in temperature can be explained by the linear relationship between mV readings and temperature readings. It can be clearly observed from the graph in figure 13 that the regression line passes almost exactly through every point on the scatter plot.

It can be clearly observed from the graph in figure 13 that the regression line passes almost exactly through every point on the scatter plot. So it can be concluded that this can be able to explain all the variations.Table 3 shows the final result of correlation coefficient which is 0.998232 and this valueindicates a strong positive linear relationship via a firm linear rule.Table 3 Correlation coefficient of the curve fitVoltage (mV) Temperature(oC)Voltage (mV) 1 0.998232Temperature(oC)14.2 Heat Conduction TestAluminum alloy 4043 which has 5% Si is used in this deposition. This alloy is used to weld many similar and dissimilar cast and wrought Al parts. It is extensively used with the GTAW and GMAW processes with Argon as shielding gas. The original thermal conductivity of the welding wire is 163 Wm-1 K-1.Six thermocouples are used to record the temperature difference on the heat input and output side of opposite faces of the sample as shown in figure 12.

Three small holes, for three thermocouples on each side, with 3mm depth are drilled into four sides of the sample. The threefixed thermocouples T1, T2 and T3 are positioned along the heater section with a gap of 16mm in between them. The other thermocouples T4, T5 and T6 are attached to the output side at similar intervals. The thermocouples on the input and output sides were connected to the selector switch for sequential selection and recording of the voltage readings.Figure 12 Schematic diagram of experimental setupThe apparatus consist of heating elements fitted on one side of the sample whose thermal conductivity has to be measured. The heater plate assembly is made up of mica sheets on which the Nichrome sheet is sandwiched in between upper and lower mica sheets for stabilizing and avoiding damage of the thermocouples. The sample was insulated on all of its faces for preventing the heat loss to the environment.

The sample was insulated on all of its faces for preventing the heat loss to the environment. When power is supplied to the heater, heat transfer will take place from the heat input side to the output side. The temperature of the sample is measured at the six positions by using a K type thermocouples and digital voltmeter. The power supply is adjusted to desired quantity by means of electronic controlled circuit by adjusting the voltage and current so as to provide the required watts of input to the heater plate. Electric power input can be measured by using digital voltmeter and ammeter multiplier [12]. The steady state temperature was ensured by the change in the voltage reading for each thermocouple which is fitted at the surface of the specimen. Steady state is reached when these voltmeter readings stop changing with time.The rectangular sample was surrounded by a glass wool with a very high insulating capability as shown in figure 13(a).

Steady state is reached when these voltmeter readings stop changing with time.The rectangular sample was surrounded by a glass wool with a very high insulating capability as shown in figure 13(a). A glass tape was used to tightly hold the sample with the heating assembly and the insulating material.

Temperature differences are measured by the calibrated thermocouples which are attached at the heat input and output sides of the specimen.(a) Insulation of sample(b) Power supply and selector switch(c) Heating plate assembly(d) Position of thermocouplesFigure 13 Sample and heat assembly for conductivity testingVariation of input and output temperature along air gaps with respect to time is shown in figure 14, it shows that almost constant temperature difference maintained between heat input and output side.Figure 14 Variation of temperature along air gaps with timeVariation of input and output temperature across air gaps with respect to time is shown in figure 15, it shows that almost constant temperature difference maintained between heat input and output side.Figure 15 Variation of temperature across air gaps with timeAs explained in the previous sections the main objective of this experiment is to prove the capability of HLM in producing directionally conductive material and to measure the directional conductivity of the sample in two different directions.

It was assumed that the small variation in the air gap width due to spattering of molten metal to adjacent weld beads is negligible. For the measurement of thermal conductivity (K) one dimensional heat flow through the rectangular specimen was considered. Fourier’s law of heat conduction was considered throughout the heat conduction test.4.2.2 Experimental ResultsThe experimental results for the DHC specimen showed anisotropic heat conduction measured through the opposite sides. A heat input of 30.4 watt was provided from one side of the specimen and the temperature difference across the specimen was measured. The input and output measured temperatures were averaged after steady state condition was reached. The variation of average temperature with time along and across air gaps shows a similar trend.

The input and output measured temperatures were averaged after steady state condition was reached. The variation of average temperature with time along and across air gaps shows a similar trend. Variation of temperature along and across air gaps with time is shown in figure 16.It shows output average temperature along air gap is more than that of temperature across air gap, which implies thermal conductivity across gap is less than that of across air gap. Where: Q= heat transfer rate = V.I =30.4 Watt, A= Area normal to heat transfer, m2 Th = temperature of hot plate i.e. (T1 + T2 + T3) /3 in °C Tc = temperature of cold plate i.e. (T4+T5 + T6) /3 in °C L = thickness of specimen along heat flow in mTo check the uni-directional heat flow, the temperature was measured on the adjacent side ofthe heat input side and a negligible heat flow was observed.The measured value of thermal conductivity was 0.1398Wm-1 0C-1, which is negligible whencompared to heat flow on the heat input side.

The variation of temperature along adjacent side of heat input is shown in figure 17, it shows there is negligible temperature difference between heat input and adjacent side.Figure 16 Variation of temperature along and across air gaps with timeFigure 17 Variation of temperature along adjacent side of heat input5. ConclusionUnlike other heat sink, which radiates from all surfaces and in all directions, the new object developed through HLM radiate elaborately in one direction. The Al-Si alloy transfer heat away from the source along the welding direction but only minimally in a direction transverse to the welding direction. The object has demonstrated heat transfer of 130 W/m-k measured parallel to air gaps and 100 W/m-k perpendicular to air gaps. In general, the value of heat conductivity parallel and perpendicular to air gaps on the specimen was showed anisotropic thermal conductivity.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAdditive Manufacturing of Energetic MaterialsM. Chiroli*, F. Ciszek*, and B. Baschung*French-German Research Institute of Saint-Louis, 5 rue du Général Cassagnou, 68301 Saint- Louis CEDEX, FRANCEAbstractOver the last decade, additive manufacturing gained a lot of interest within the defense industry. However, printing energetic materials such as explosives or solid propellants remains a challenge. A research work with the aim to suggest appropriate materials and associated 3D printing techniques to produce energetic materials, is currently carried out. In interior ballistics, improved weapon performance ends in reaching higher muzzle kinetic energy. This can be achieved by adjusting the gas pressure released during the ballistic cycle by using specific propellant grains.

This can be achieved by adjusting the gas pressure released during the ballistic cycle by using specific propellant grains. A former study (Baschung, MRS Symposium H, Boston, 2005) highlighted the advantages of using co-layered propellant grains to achieve this goal. The difficulties in producing such kind of propellants can now be bypassed thanks to 3D printers. Our approach consists in deposing a viscous energetic paste, containing a solvent, with an adapted machine. The layer formation and the adhesion between them are investigated in this work.1. IntroductionIn recent years, additive manufacturing (AM) technologies have experienced considerable growth. The various existing techniques enable to produce parts with different materials, such as metal alloys, polymers, ceramics or even concrete [1] [2] [3] [4]. Thus, for example, 3D printing techniques can now be used to produce airfoils for aerospace applications, automotive parts or medical implants [5].

Thus, for example, 3D printing techniques can now be used to produce airfoils for aerospace applications, automotive parts or medical implants [5]. This wide range of possibilities led the defense sector to take a close interest in this process. For this reason, an increasing number of studies are currently carried out to use additive manufacturing within the defense industries [6].An energetic material (EM) is a chemical substance that can release its stored energy in a very short time. Once initiated, these materials undergo self-sustaining exothermic reactions. They can be classified as explosives, propellants and pyrotechnics. The latter can be used in countermeasure decoy flares for the safety of military aircrafts [7] or in fireworks. Propellants can be used in airbag systems [8] and in space launchers [9]. They can also be used in interior ballistics which is the branch of applied physics dealing with the motion of the projectile in the gun until it exits the barrel.

They can also be used in interior ballistics which is the branch of applied physics dealing with the motion of the projectile in the gun until it exits the barrel. The final performance of energetic materials is linked to their chemical composition and geometry.The main current issue within the field of interior ballistics consists in increasing the weapon performances. To achieve this goal, it is necessary to raise the muzzle kinetic energy of the projectile propelled. One of the most intuitive ways consists in adjusting and controlling the combustion gases released from the energetic material during the ballistic cycle. This can be achieved thanks to the use of specific solid propellant grains such as multi-material propellants (co-layered propellants) [10].

This can be achieved thanks to the use of specific solid propellant grains such as multi-material propellants (co-layered propellants) [10]. Quite difficult to manufacture before, it appears that1003the production of this type of solid propellant can now be easier to handle thanks to 3D printing technologies.The use of inert materials combined with slightly modified machines can be of interest to print energetic materials. Indeed, it is possible to adapt material extrusion based machines to print highly viscous materials. For example, Yang et al. [11] mounted a jet valve on a Fab@Home v2 3D printer while Hong et al. [12] linked a screw extruder to feed the extruder of the printer. In order to produce ceramic parts, Faes et al. [13] developed and tested an additive manufacturing process based on the extrusion of a composition which has the capacity to cross- link under UV light.

[13] developed and tested an additive manufacturing process based on the extrusion of a composition which has the capacity to cross- link under UV light. Recently, scientists from the Massachussets Institute of Technology (MIT) investigated the additive manufacturing of cellulosic materials. Pattinson and Hart [14] succeeded in developing a simple process, operating at room temperature, for the 3D printing of cellulose. They dissolved cellulose acetate (CA) powder in acetone and modified a commercial-off-the-shelf fused deposition modeling (FDM™) machine (Printrbot Simple Metal) to meet their objectives. Studies directly concerning energetic materials are currently emerging. Indeed, Fleck et al. [15] prepared a reactive filament made of aluminum (Al) and polyvinylidene fluoride (PVDF) that can be printed on an unmodified FDM™ 3D printer. The Dutch research center TNO started working on AM of energetic materials in 2013 [16].

The Dutch research center TNO started working on AM of energetic materials in 2013 [16]. After having successfully printed trinitrotoluene (TNT) with a FDM™ printer, because of the expected performance gain, TNO focused on the 3D printing of gun propellants. They succeeded in printing multi-perforated gun propellants containing up to 75 wt% of energetic material [17] by using a stereolithography based printer.The objective of this study is to suggest appropriate materials and associated 3D printing techniques to produce energetic materials, specifically gun propellants. The idea, based on the work of Pattinson and Hart [14], is to use viscous pastes that can be printed on a slightly modified material extrusion based 3D printer. Materials (both inert and energetic) are selected to prepare viscous mixtures ready to be deposited.

Materials (both inert and energetic) are selected to prepare viscous mixtures ready to be deposited. The approach developed consists in using an inert paste, equivalent to the energetic one, to study the layer formation process and to configure the 3D printer. Different inert compositions in terms of viscosity are prepared. Dynamic viscosity measurements are carried out and the behavior of the pastes during extrusion is analyzed. To understand the layer formation process, deposition and adhesion tests are performed. Once the best inert candidate has been identified and the printer has been configured accordingly, an energetic paste having the same viscosity has to be printed.2. Materials and Experimental setupTo prepare viscous pastes, seven different inert materials were hermetically kept in a solvent for 24 hours (see figure 1). Because of its capacity to reliably dissolve solid propellants, the solvent chosen was acetone (Zundel + Kohler).

Because of its capacity to reliably dissolve solid propellants, the solvent chosen was acetone (Zundel + Kohler). The considered materials are: ethylene vinyl acetate (EVA) LEVAPREN® 700 HV (Bayer AG); cellulose acetate (CA) (Sigma Aldrich); ethylene vinyl alcohol EVAL™ (Kuraray); polyninyl alcohol Rhodoviol 4/125 (Prolabo); Ultramid® 6A (BASF); polylactic acid (PLA) (Printrbot) and acrylonitrile butadiene styrene (ABS) (Ultimaker).1004Figure 1 : Solubility tests of inert materials in acetone. From left to right: EVA LEVAPREN® 700 HV (1); CA (2); EVAL™ (3); Rhodoviol 4/125 (4); Ultramid® 6A (5); PLA (6) and ABS (7)Among the inert materials tested, only two were soluble in acetone: EVA LEVAPREN® 700 HV and CA. Thus, at first, inert pastes were prepared by dissolving 10 g of ethylene vinyl acetate in 6.4 mL of acetone, and 10 g of cellulose acetate in 25.5 mL of acetone. The mixtures were manually mixed and hermetically sealed for 24 hours.

The mixtures were manually mixed and hermetically sealed for 24 hours. Feedstock was then loaded into 30 mL syringes (Nordson EFD) and extrusion were performed thanks a Performus™ V dispenser (Nordson EFD).The EM considered for this study is an in-house developed formulation. It is made of 54 wt% of nitrocellulose, mechanically mixed with other energetic ingredients. This solid propellant is soluble in acetone, and, depending on the weight percentage, different viscosities of the mixture can be reached. In the same way as it was done for inert pastes, a 30 mL syringe was loaded with an energetic paste initially prepared by dissolving 10 g of solid propellant in 10 mL of acetone. Extrusion tests were performed as well.

Extrusion tests were performed as well. For both inert and energetic pastes, dynamic viscosity measurements were done thanks to a RM 100 Plus viscometer and its MS- R4 measuring system from Lamy Rheology.The 3D printer used in this study is a material extrusion based desktop printer (Printrbot Simple Metal). This FDM™ printer has been enhanced by replacing the original filament extruder by the Paste & Food Extruder system from Printrbot. This system enables to use viscous materials loaded into syringes to 3D print parts. Initially designed to use HSW 50 mL syringes, the system has been slightly modified to use 55 mL syringes from Nordson EFD. Furthermore, for pyrotechnic safety reasons, the aluminum parts of the printer are currently being replaced by plastic ones. In case of non-desired reaction of the energetic loading, the plastic parts will generate less dangerous fragments and enable a better deconfinement of the associated pressure wave.3.

In case of non-desired reaction of the energetic loading, the plastic parts will generate less dangerous fragments and enable a better deconfinement of the associated pressure wave.3. Results and DiscussionThe first step consisted in performing extrusion tests with the three different compositions (EVA/Acetone, CA/Acetone and EM/Acetone). An examination of the material behavior during the extrusion was carried out. During extrusion from the syringe, EVA/Acetone composition showed significant swelling while, in the same conditions, no swelling or shrinkage was noticed with CA based mixture (see figure 2). Furthermore, the extrusion of CA paste was quite similar to the EM/Acetone paste. Indeed, both CA/Acetone and EM/Acetone filaments seemed close in terms of adhesive properties, while the EVA/Acetone filament turned1005out to be highly sticky.

Indeed, both CA/Acetone and EM/Acetone filaments seemed close in terms of adhesive properties, while the EVA/Acetone filament turned1005out to be highly sticky. Consequently, the CA/Acetone paste was retained to be used as an inert composition.Figure 2 : Extrusion tests performed with the three compositions. From left to right:EVA/Acetone; CA/Acetone; EM/Acetonefor deeper characterizations. Two different samples of cellulose acetate (sample A: molecular weight Mn ≈ 30,000 by GPC and sample B: molecular weight Mn ≈ 50,000 by GPC) were dissolved from 25 wt% up to 50 wt% in acetone by manual mixing. The details of the prepared samples are given in table 1 and 2.

The details of the prepared samples are given in table 1 and 2. An energetic composition with a charge rate of 60 wt%, looking like the more viscous inert compositions, was also prepared (see table 3).Therefore, differentinert mixtures CA/Acetone were preparedTable 1 : CA/Acetone compositions prepared with sample A (Mn ≈ 30,000)Name Weight CA (g) Volume Acetone (mL) Charge rate (wt%) A25wt% A30wt% A35wt% A40wt% A45wt% A50wt%10.01 12.86 16.16 20.04 24.55 30.0138.4 38.3 38.4 38.3 38.3 38.325 30 34.9 40 45 50Table 2 : CA/Acetone compositions prepared with sample B (Mn ≈ 50,000)Name Weight CA (g) Volume Acetone (mL) Charge rate (wt%) B25wt% B30wt% B35wt% B40wt% B45wt% B50wt%10.01 12.89 16.15 20.00 24.55 30.0038.3 38.4 38.3 38.3 38.3 38.325 30 35 40 45 50Table 3 : EM/Acetone compositionName Weight EM (g) Volume Acetone (mL) Charge rate (wt%)EM60wt%4538.360The dynamic viscosity of the prepared mixtures was measured.

As it was not possible to obtain homogenous pastes for highly charged preparations, no viscosity measurements could be conducted for samples A50wt%, B45wt% and B50wt%. The velocity gradient value retained for the measurements was determined thanks to the visually more viscous preparation1006(A45wt%). It was set to 4 s-1. Measurements were realized at room temperature (22 °C ± 2 °C). Results can be seen in figure 3. Because of the higher molecular weight of sample B of CA, for a same charge rate, compositions prepared with this sample are more viscous than sample A based preparations (see figure 3). This enables to screen a wide range of inert compositions from syrup like fluids to mastic like ones. Furthermore, it is noticed that the dynamic viscosity of the EM60wt% (1126 Pa.s at 4 s-1) composition is in the same range than the one of A45wt% (1388 Pa.s at 4 s-1) and B40wt% (1238 Pa.s at 4 s-1) pastes.

In terms of dynamic viscosity, it is thus possible to have an inert composition equivalent to the energetic pastes.Figure 3 : Viscosity measurements of CA/acetone compositionsIn order to study the layer formation process, extrusion and deposition tests were performed. Thus, the prepared pastes were extruded through different tapered tips (Nordson EFD, diameter Φ from 0.84 mm to 0.25 mm) to obtain a paste filament that can be manually deposited onto glass microscope slides, simulating the build plate. For both the energetic and inert mixtures, the extruded paste filaments were sticky when acetone was still present, but they solidified with the acetone evaporation. Kapton® covered glass slides showed better adhesion than bare glass. Indeed, when deposited on bare glass, the materials first adhere, but rapidly slide with the solvent evaporation. The pressure needed to extrude the compositions through the different types of tips was measured.

The pressure needed to extrude the compositions through the different types of tips was measured. The tests were considered as successful if it was possible to vertically deposit the extruded filament on the microscope slides. Results for the peak viscosity values as well as for an intermediate one are given in table 4. EM based paste results can be seen in table 5. A green box means that the test was successful while a red one means that it failed.1007Table 4 : Pressure needed to extrude the inert pastesPressure (bar)CA/Acetone A25wt%CA/Acetone B35wt%CA/Acetone A45wt%Φ tip (mm) 0.84 0.58 0.41 0.25 0.84 0.58 0.41 0.25 0.84 0.58 0.41 0.25Φ tip (mm)Φ tip (mm)2 3 4 5 6 7Table 5 : Pressure needed to extrude the energetic pasteEM/Acetone EM60wt%Pressure (bar)Φ tip (mm) 0.84 0.58 0.41 0.252 3 4 5 6 7As shown in table 4 and 5, a similar extrusion behavior can be noticed for the EM60wt% and the A45wt% compositions.

This reinforces the fact that the CA/Acetone composition can be considered as the inert equivalent to the EM/Acetone paste. To obtain the best surface finishes, the thinner tip (0.25 mm) should be used for the paste extrusion, hence the need of being able to supply 5 bar to extrude the more viscous compositions (see table 5). The pressure needed corresponds to a force of around 300 N applied to the piston of the syringe. This threshold can easily be reached with the Paste & Food Extruder system.Printing tests were performed with the CA/Acetone compositions. They were conducted at room temperature and no part was heated on the printer (build plate or extruder). For safety reasons, no printing test was carried out with the EM/Acetone mixture until now because prior authorization must be obtained for using EM. The printing parameters used for the first tests are the ones used to print PLA material. No optimization of these parameters was attempted so far.

The printing parameters used for the first tests are the ones used to print PLA material. No optimization of these parameters was attempted so far. The tests consisted in printing the contours of the first layers of a 3 mm side calibration box. For each inert composition, it was not possible to build a perfect and reliable first layer (see figure 4). This should be solved by optimizing the printing parameters. However, those first attempts seemed encouraging, showing the possibility to reach a good adhesion of the paste on the build plate. Moreover, the hardening of the deposited composition turned out to be in the same time range as compared to standard 3D printing materials. Attempts to build successive layers were made.

Moreover, the hardening of the deposited composition turned out to be in the same time range as compared to standard 3D printing materials. Attempts to build successive layers were made. In order to receive another layer, it seemed that one layer should be completely dried and solidified.1008Verygood adhesionof the first depositedlayerImpossible to builda perfectfirst layer withthe printing parametersusedFigure 4 : First printing tests performed with CA/Acetone paste (contours of a 3 mm box)4. Conclusion and Future workIn this paper, two different kinds of viscous pastes were prepared with the aim to be printed with a material extrusion based desktop printer. The compositions were obtained by dissolving either an energetic material (EM) or cellulose acetate (CA) in acetone. The CA/Acetone composition was selected as an inert equivalent to the EM/Acetone mixture and should be used to configure the 3D printer.

The CA/Acetone composition was selected as an inert equivalent to the EM/Acetone mixture and should be used to configure the 3D printer. Dynamic viscosity measurements were performed to identify an equivalence between inert and energetic compositions. Extrusion and deposition tests highlighted the similar behavior of both inert and energetic pastes. Specifically, it was found that inert mixtures (A45wt% and B40wt%) resemble the properties of an energetic counterpart (EM60wt%) in terms of dynamic viscosity. Thus, testing and configuring the setup can be conducted with the much safer inert paste. Eventually, first printing experiments carried out with the inert compositions turned out to be encouraging. Furthermore, the necessity to have one layer completely dried and solidified before receiving the next one was noticed.The next steps of this research work are, first, to optimize the printing of the inert mixture and to identify the best printable inert candidate.

Second, an energetic paste, within the same range of dynamic viscosity than the previously identified inert candidate, should be prepared. A first simple shape of solid propellant (e.g. flake) should eventually be printed and its performances evaluated.5. AcknowledgementsThe authors greatly thanks the French state organization responsible for armament programs, the DGA, for supporting this research. The thanks are also extended to S. Braun, J. Glorian, K.-T. Han and R. Wölbing for their support and advice.10096. References[1] K. V. Wong and A. Hernandez, "A Review of Additive Manufacturing," ISRNMechanical Engineering, 2012.[2] S. C. Ligon, R. Liska, J. Stampfl, M. Gurr and R. Mülhaupt, "Polymers for 3D Printing and Customized Additive Manufacturing," Chemical Reviews, vol. 117, no. 15, pp. 10212-10290, 2017.[3] T. D. Ngo, A. Kashani, G. Imbalzano, K. T. Nguyen and D.

Mülhaupt, "Polymers for 3D Printing and Customized Additive Manufacturing," Chemical Reviews, vol. 117, no. 15, pp. 10212-10290, 2017.[3] T. D. Ngo, A. Kashani, G. Imbalzano, K. T. Nguyen and D. Hui, "Additive manufacturing (3D printing): A review of materials, methods, applications and challenges," Composites Part B, vol. 143, pp. 172-196, 2018.[4] Y. L. Yap, V. Dikshit, S. P. Lionar, H. Yang, J. C. Lim, X. Qi, W. Y. Yeong and J. Wei, "Investigation of Fiber Reinforced Composite using Multi-Material 3D Printing," in 27th Solid Freeform Fabrication Symposium, Austin, 2016.[5] N. Guo and M. C. Leu, "Additive manufacturing: technology, applications and researchneeds," Frontiers of Mechanical Engineering, vol. 8, no. 3, pp. 215-243, 2013.[6] A. Busachi, J. Erkoyuncu, P. Colegrove, F. Martina, C. Watts and R. Drake, "A review of Additive Manufacturing technology and Cost Estimation techniques for the defence sector," CIRP Journal of Manufacturing Science and Technology, vol.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAdditive Manufacturing of Fatigue Resistant Materials: Avoiding the EarlyLife Crack Initiation Mechanisms during FabricationJonathan Pegues1,2, Michael Roach3, Nima Shamsaei1,2*1Department of Mechanical Engineering, Auburn University, Auburn, AL 2National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 3Department of Biomedical Materials Science, University of Mississippi Medical Center, Jackson, MSCorresponding author: Email: shamsaei@auburn.edu Phone: (334) 844 4839AbstractThe full potential of additive manufacturing (AM) has been limited by the process induced defects within the fabricated materials.

Defects such as lack of fusions and gas entrapped pores act as stress concentrators and result in premature fatigue crack initiation, severely limiting the applicability of AM in fatigue-critical applications. However, by understanding the failure mechanisms associated with AM materials and leveraging the intimate localized thermal input (i.e. process conditions), the failure mechanisms for some materials may be avoided during the fabrication process. This study investigates the crack initiation behavior of an AM austenitic stainless steels subjected to fatigue testing. The microstructural features responsible for fatigue crack initiation are captured at the surface by ex-situ electron backscatter diffraction.

The microstructural features responsible for fatigue crack initiation are captured at the surface by ex-situ electron backscatter diffraction. Results show that the higher cooling rates during AM offer the opportunity to fabricate fatigue resistant austenitic stainless steel parts by avoiding the microstructural features that are most detrimental to fatigue performance.Keywords: Laser powder bed fusion, Crack nucleation, Twin boundary, Heat treatment542IntroductionAdoption of additive manufacturing for fatigue critical applications has been plagued by the presence of process induced defects such as surface roughness and/or porosity. These defects introduce significant stress concentrations leading to early fatigue life crack initiation, which can significantly affect the fatigue behavior of additive manufactured parts.

These defects introduce significant stress concentrations leading to early fatigue life crack initiation, which can significantly affect the fatigue behavior of additive manufactured parts. While most researchers have focused on optimizing the additive manufacturing process to achieve maximum density, the fabrication of defect free AM parts has not been realized.Researchers have long suggested that establishing the feedstock-process-structure- property-performance (FPSPP) relationships could lead to the ability to additively manufacture more fatigue resistant materials [1]. The goal of applying this approach is to leverage the material and processes effects on the microstructure to tailor the material properties for a specific service condition and ultimately improve part performance.

This type of approach is essentially application driven, tailoring of the microstructure and is accomplished through the intimate relationship between the supplied thermal history and the resulting microstructure. To date, however, this approach has been overshadowed by the effect of defects, which typically outweigh any improvements in microstructure [1-3].Austenitic stainless steels are exceptionally tough alloys having moderate strength and high ductility. Unlike most engineering alloys, austenitic SSs are less susceptible to inclusions and porosity that typically plague the high cycle fatigue (HCF) behavior of common metallic materials [4, 5]. Inclusions and porosity, create stress concentrations that often lead to crack initiation and ultimately fatigue failure.

Inclusions and porosity, create stress concentrations that often lead to crack initiation and ultimately fatigue failure. For austenitic SSs, with exceptional toughness, can accommodate these types of stress concentrations without initiating early life fatigue cracks and are instead sensitive to microstructural features [4, 6].Previous studies have revealed that annealing twin boundaries (Σ 3-TB) within the microstructure of 304L and 316L SSs are the most susceptible to crack initiation [4, 5, 7-9]. Specifically, Σ 3-TB that are oriented on the surface at approximately 45° from the loading direction are the most likely feature to initiate a dominant crack leading to failure. The deformation characteristics of these boundaries results in elastic incompatibilities near the boundary which lead to compliance stresses which can result in slip behavior near the boundary. As slip accumulates, persistent slip bands form resulting in intrusions/extrusions near the boundaries.

As slip accumulates, persistent slip bands form resulting in intrusions/extrusions near the boundaries. These intrusions/extrusions are known to create micro-cracks that can develop into propagating cracks not only at Σ 3-TB but high angle grain boundaries (HAGB) as well [4]. This same crack initiation behavior can be exacerbated under increasingly greater loads which can create plastic incompatibilities in conjunction with the elastic ones leading to much faster crack formation.It was suggested previously that the ability to control the Σ 3-TB/HAGB orientations or density during fabrication could lead to improved fatigue resistance [4], however, the means to control these specific boundaries was not known at the time. The emergence of additive manufacturing as an advanced manufacturing technique, however, has supplied the catalyst to manufacture austenitic stainless steel with limited Σ 3-TB/HAGB.

The emergence of additive manufacturing as an advanced manufacturing technique, however, has supplied the catalyst to manufacture austenitic stainless steel with limited Σ 3-TB/HAGB. In this study, the fatigue behavior of laser beam powder bed fusion (LB-PBF) austenitic stainless steels is investigated. These stainless steel alloys may provide early insight into how the FPSPP relationships can be543leveraged to fabricate more fatigue resistant materials by avoiding the typical failure mechanisms responsible for crack initiation.Experimental MethodsRectangular 304L SS specimen blanks were fabricated horizontally on an EOS M290 using standard 316L stainless steel process parameters. Specimens were stress relieved at 400°C for 1 hr and furnace cooled before being removed from the substrate. The blocks were then machined into standard cylindrical dogbone fatigue (CDF) specimens and non-standard square gage fatigue (SGF) specimens as shown in Figure 1.

The blocks were then machined into standard cylindrical dogbone fatigue (CDF) specimens and non-standard square gage fatigue (SGF) specimens as shown in Figure 1. A set of specimens from each alloy was solution annealed at 1050°C for 1 hr then water quenched. All specimens were longitudinally polished to a mirror finish using silicon carbide grinding papers then further electro-polished using Struers A2 solution to obtain a pristine surface for EBSD analysis.a)I112.80 r R25 C=i b 1:j I .,..I ·~I33.75I70.01Ib)145 .29R 101.60I • ,asGage W idthFigure 1: Specimen geometry for (a) cylindrical dogbone and (b) square gage fatigue specimens.Microstructural characterization was carried out on a Zeiss FE-SEM using an EDAX electron backscatter diffraction (EBSD) detector and OIM software. Analysis was performed on each surface.

Analysis was performed on each surface. Surfaces whose normal is perpendicular to the build direction are defined as ZXY and surfaces with normal is parallel with the building direction are defined as YXZ per ASTM WK49949 [10]. These directions, in relation to the specimen, are shown in Figure 2. All fatigue tests were conducted on an MTS servo-hydraulic test frame with 100 kN capacity. The CDF specimens were tested at several stress amplitudes until complete failure while the SGF specimens were tested at a single stress amplitude and interrupted starting at 10% of their expected life to failure to observe crack initiation. During the interrupted periods the specimens were loaded into the SEM and micro cracks were located and mapped using EBSD.

During the interrupted periods the specimens were loaded into the SEM and micro cracks were located and mapped using EBSD. A constant loading rate was used for all fatigue testing to avoid any load rate effects for these materials.544Z - Build DirectionFigure 2: Orientation and directions for the square gage fatigue specimens.Experimental ResultsMicrostructureThe microstructural characteristics for both (a) 304L SS stress relieved (304L-SR) and (b) 304L SS solution annealed (304L-SA) conditions are similar and are shown in Figure 3. On the left is the inverse pole figure (IPF) plot with the corresponding grain boundary misorientation chart and pole figure to the right. There are some remarkable characteristics for both alloys including a <011> texture, low density of HAGBs, and no easily-identifiable Σ 3-TB. The <011> texture, which is approximately 4X the random, occurs due to the <001> easy growth direction and the fast moving heat source. Sun et al.

The <011> texture, which is approximately 4X the random, occurs due to the <001> easy growth direction and the fast moving heat source. Sun et al. [11] showed that scanning strategy can result in various melt pool geometries with shallow melt pools showing a <001> texture and deeper melt pools resulting in the <011> texture observed in this study. It should be noted, however, the texture is not perfectly aligned with the machined surface of the specimen as indicated by the 011 pole figure being slightly rotated about A2.545a)A10.50.4C: 0 :.:; ue ~ ., .0 0.2 E :::, z0.3lL0 1 10.

1102030405060Misorientation Anqle [degrees)A1b)0.5C: 0 :.:;0.4A2u e 0.3 :;; .0 0.2 E :::, zlL0 1 10.10.0102030405060Misorientation Angle [degrees)Figure 3: Inverse pole figure plot with grain boundary misorientation plots and pole figures for (a) 304L-SR and (b) 304L-SA conditions.The grain morphology for YXZ and ZXY surfaces were observably different for both 304L-SR and 304L-SA specimens as shown in Figure 4. The ZXY surfaces grains were slightly elongated in the build direction where the YXZ grains were mostly equiaxed in its plane. Table 2 shows the average grain diameters and standard deviations for both surfaces and for both thermal treatment condition. Due to the elongated grains in the build direction, the ZXY surfaces had larger average grain diameters and standard deviations than the YXZ surfaces. The 304L-SA specimens showed a roughly 12% increase in grain size for both YXZ and ZXY surfaces.

The 304L-SA specimens showed a roughly 12% increase in grain size for both YXZ and ZXY surfaces. The differences in grain sizes for each surface, though small, can result in different deformation behaviors for each surface.

The differences in grain sizes for each surface, though small, can result in different deformation behaviors for each surface. For instance, larger grains typically show higher slip activity than smaller grains which can have significant implications on the crack initiation behavior as it is typically dominated by dislocation activity at the surface [12, 13].546Stress RelievedSolution AnnealedFigure 4: Micrographs for 304L-SR (top) and 304L-SA (bottom) showing grain structure for YXZ surfaces (left) and ZXY surfaces (right).Table 1: Grain sizes for 304L-SR and 304L-SA for both YXZ and ZXY surfaces.SurfaceGrain Size (μ m)Standard Deviation (μ m)Stress Relief – 400°C 1-hr, furnace cooledYXZ11.948.75ZXY13.2810.12Solution Anneal – 1050°C 1-hr, water quenchYXZ13.4410.48ZXY14.7512.13547Fatigue BehaviorThe fatigue results, surprisingly, show improved fatigue resistance for the 304L-SR condition compared to wrought 304L stainless steel.

Figure 5 shows the stress-life fatigue results and corresponding Basquin curves. The 304L-SR condition showed an increase in fatigue life one to two orders of magnitude for a given stress amplitude. This improved fatigue resistance is surprising for LB-PBF material, especially in the high cycle fatigue (HCF) regime which is often dominated by crack initiation at defects. The majority of the HCF life is often spent in the crack initiation stage which suggest that the improved fatigue resistance is a result of improved resistance to crack initiation.500,,-.._ ro ~ :E _, bro 400 <1.f '"O ;::s -- ...... ...... 0.. ~ 300¢..~□ - -l -- - -- -ITT04L-SR ◊ 304L-Wroughtrn- .... " - ~-- --~~ .... _ .... -0.

...... 0.. ~ 300¢..~□ - -l -- - -- -ITT04L-SR ◊ 304L-Wroughtrn- .... " - ~-- --~~ .... _ .... -0. -□ ---- □I----□rJ:J rJ:J ~1-, --VJ.2001E+031E+04lE+0S Cycles to Failure, Nf1E+061E+07Figure 5: Stress-life fatigue data and related Basquin curvesThe ex-situ crack initiation investigation was carried out on the SGF specimens to better understand the failure mechanism associated with the LB-PBF material. Figure 6 shows cracks at 10% of the fatigue life for the stress amplitude of 330 MPa. The 304L-SR specimens showed similar crack lengths at 10% of the HCF life compared to the 304L SS wrought (304L-W) counterpart. Despite the high number of fatigue cycles, the 304L-SR specimens show far less slip activity at the surface as indicated by the persistent slip bands (PSB) in the SEM micro-graphs. Additionally, the 304L-W material shows more developed cracks than the 304L-SR material at a fifth of the fatigue life.

Additionally, the 304L-W material shows more developed cracks than the 304L-SR material at a fifth of the fatigue life. Each crack in the 304L-W condition appears to have been deflected by grain boundaries resulting in roughness induced closure and keeping the middle portion of the548crack open. On the other hand, the 304L-SR material shows mostly straight cracks that appear to not have been deflected by neighboring grains.304L-SR304L-Wroughtj60µmI20µm'I 20,000 Cycles30 µm.. ., . t.20, ooo C cresFigure 6: Observed cracks in 304L-SR (left) and 304L-W (right).The EBSD maps were collected for several cracks to understand mechanisms governing the crack initiation behavior. In order to better understand how the cracks are influenced by the surrounding microstructure, the 304L-SR SGF specimens were cycled further to 300,000 cycles or approximately 30% of the expected fatigue life.

Figure 7 shows the IPF map on the left and the Taylor factor map on the right for both 304L-SR cracks shown in Figure 6. Despite being cycled up to 30% of their expected fatigue lives both cracks showed minimal propagation into the surrounding microstructure. This is in stark contrast to the wrought material which after 35,000 cycles had developed a dominant crack approximately 300 μ m in length. Unlike the wrought549material where mismatches in Taylor factors across the grain boundary resulted in elastic incompatibilities and crack nucleation [4], the cracks in the 304L-SR material are occurring at grain boundaries with similar Taylor factors.

Interestingly, both cracks developed in a region with higher Taylor factors and are bounded by regions of lower Taylor factors.Figure 7: IPF maps (left) and Taylor factor maps (right) for the two 304L-SR cracks shown in Figure 6.The increase in fatigue strength for LB-PBF 304L SS appears to be attributed to a combination of differences in the as-fabricated microstructure. Most importantly the reduced density of Σ 3-TB and HAGB in the 304L-SR material results in much fewer optimal crack initiation sites for this material. Secondly, the finer microstructure restricts dislocation movement as evident by comparing the 304L-SR and 304L-W SEM images in Figure 6. The wrought material with much larger grains shows a much higher density of PSB on the surface leading to fewer intrusion/extrusions which are also prime crack nucleation features.

The wrought material with much larger grains shows a much higher density of PSB on the surface leading to fewer intrusion/extrusions which are also prime crack nucleation features. The PSBs can be observed in relatively large grains with higher Taylor factors, while they are absent from smaller grains that550have low Taylor factors and more efficient slip systems. This indicates that grain size plays a greater role in crack initiation than crystal orientation for 304L-SR.The effect of grain size on the crack initiation behavior was investigated further using the 304L-SA material which had larger grains as a result of the high temperature annealing procedure. Figure 8 shows the IPF and Taylor factor map of a fatigue crack that has developed in the 304L- SA specimen after only 20,000 cycles.

Figure 8 shows the IPF and Taylor factor map of a fatigue crack that has developed in the 304L- SA specimen after only 20,000 cycles. Other than the larger grain size, it can be observed in the Taylor factor map that the crack has occurred at a grain boundary that is approximately 45° from the loading direction and has a significant mismatch in Taylor factor across the HAGB. In fact, all cracks that were observed in the 304L-SA material occurred either at Σ 3-TB/HAGB with mismatched Taylor factors or at discontinuities such as pores on the surface in the vicinity of grains with large Taylor factor mismatches.

This observation indicates that the larger grain size results in a shift in the crack initiation mechanism back in line with what is typically observed in 304L-W material.Figure 8: Representative IPF map (left) and Taylor factor map (right) for a crack observed in the 304L-SA conditions.ConclusionsThe aim of this study was to better understand the structure-property relationships of LB- PBF 304L SS. Specifically, the failure mechanisms associated with fatigue crack initiation was investigated through ex-situ experimental EBSD observations of crack initiation and the features resulting in crack initiation. Results indicate that the as-fabricated microstructure in the stress relieved condition has a much higher fatigue strength related to the avoidance of the typical crack initiation mechanisms for wrought 304L SS.

The fine microstructure and low Σ 3-TB/HAGB density of the 304L-SR material leads to limited preferred crack initiation sites with cracks forming mainly at HAGB separating grains that are much larger than the average grain size. These cracks propagate until they reach smaller grains after which they slow significantly. This change in crack initiation mechanisms directly leads to the higher strength observed for the 304L-SR condition.551Furthermore, solution annealing the LB-PBF material results in some grain growth with the average grain diameter increasing by approximately 12% from the average diameter of the stress relieved condition. Observations of the crack initiation behavior for the 304L-SA condition showed a shift in the mechanisms back to what is typically observed for wrought 304L SS.

Observations of the crack initiation behavior for the 304L-SA condition showed a shift in the mechanisms back to what is typically observed for wrought 304L SS. Mismatches in Taylor factors along HAGB, leading to elastic incompatibilities and higher localized stress states, were the most frequently observed causes of crack initiation in the 304L- SA material. This resulted in crack formation much earlier in the fatigue life of the 304L-SA condition compared to the 304L-SR material. Thus, by better understanding the feedstock-process- structure-property relationships, additive manufacturing can be used to fabricate net-shaped/near net-shaped parts with improved fatigue resistance.AcknowledgementsPartial funding for this work was provided by the National Science Foundation under Grant No. 1657195.References[1] N. Shamsaei, A. Yadollahi, L. Bian, S.M.

Composites of PPS and PEKK are emerging as viable candidates for several components in aerospace and tooling industries and additive manufacturing of these materials can be extremely beneficial to lower manufacturing costs and lead times. However, these materials pose several challenges for extrusion and deposition due to some of their inherent properties as well as thermal and oxidative responses. To better understand the properties of such systems specific to 3D printing and determine the critical parameters that make them “printable”, various rheological and thermal properties have been studied for neat as well as short fiber reinforced PPS and PEKK systems.

Attempts were also made to print these materials in a customized high temperature fused filament fabrication system.This workinvestigatesthe use ofIntroductionHigh performance thermoplastics and their fiber reinforced composites are widely used in advanced applications that demand properties such as thermal stability, high continuous use temperature, chemical, wear and flame resistance, and superior mechanical properties. Some of the commonly used high performance or high temperature thermoplastics include amorphous thermoplastics such as polyetherimide (PEI), polyphenyl sulfone (PPSU), and polyether sulfone (PES) and semi-crystalline thermoplastics such as polyphenylene sulfide (PPS), poly(etheretherketone) (PEEK), and poly(etherketoneketone) (PEKK) [1].

Semi- crystalline thermoplastics offer certain advantages over amorphous thermoplastics in terms of higher range of continuous use temperature, the ability to use the material at temperatures even above Tg for short spans without much loss in modulus [2], better resistance to creep deformation [2], and excellent chemical and wear resistance. Many of these properties can be modified by varying the degree of crystallinity.Specifically, for additive manufacturing (AM) applications, the inherent property of these materials to crystallize offers the potential of additional bonding mechanisms between the printed layers. In amorphous polymers, thermal fusion and polymer interdiffusion is the primary bonding mechanism between printed layers [3]. However, in the case of semi- crystalline polymers, interlayer bonding can be enhanced by mechanisms such as co- crystallization.

However, in the case of semi- crystalline polymers, interlayer bonding can be enhanced by mechanisms such as co- crystallization. In some studies done by Gauthier Jarrousse on self-adhesion of semi-crystalline polymers, it was shown that co-crystallization can be an effective method to reinforce interfaces if preceded by some interdiffusion of chains prior to recrystallization [4]. It has also been shown that for semi-crystalline FDM filaments, having a core-shell structure with906different peak crystallization temperatures can reduce the level of curling and distortion and also improve interlayer strength [5].Filaments for fused filament fabrication (FFF) or fused deposition modeling (FDM) processes have been produced using high temperature semi-crystalline materials in recent years.

INDMATECH GmbH has developed raw PEEK FDM filaments [6], 3DXTech has developed FirewireTM carbon fiber reinforced PEEK filaments printable at 360oC to 390oC [7] and PPS filaments that can be FDM printed between 3250C and 3450C [8]. Stratasys has developed electrostatic discharge ESD PEKK filaments that have conductive fillers in PEKK base resin [9].The goal of this study is to explore the use of neat as well as carbon fiber reinforced grades of PPS and PEKK as potential FFF feedstock materials. Thermal analysis of these materials is initially used to identify processing temperature range, followed by rheological analysis to study the effect of parameters such as temperature, shear, and fillers on the viscosity of the selected grades.

Upon identifying suitable processing conditions, filaments were made with the PPS and PEKK grades and extrusion was attempted on a customized desktop scale FFF system to determine if the materials would extrude without clogging.ExperimentalThree grades each of PPS and PEKK were procured in the form of pellets. Neat PPS and PPS filled with 40 wt. % and 50 wt. % short carbon fiber were supplied by Techmer Engineered Solutions. Three grades of PEKK from the Kepstan® 8000 series - neat PEKK and PEKK filled with 30 wt.% and 40 wt.% short carbon fiber were supplied by Arkema Inc. Prior to all tests, all PPS pellets were dried at 135oC for 2 hours and all PEKK pellets were dried at 150oC for about 3 hours.Thermal AnalysisThe first step is to determine the temperature range in which the materials would be processable in an extrusion based system.

To identify the upper limit or the decomposition temperature (above which volatiles are released), thermogravimetric (TGA) analysis was conducted using a TA Instruments Q500 instrument. The dried pellets were heated at a rate of 10oC/min in air from 25oC to 800oC.The lower limit of processing temperature is the melting temperatures, which was determined by Differential Scanning Calorimetry (DSC) using a TA Instruments Q2000 system. The dried pellets were heated at the rate of 10oC/min from 25oC up to 325oC for PPS grades and up to 370oC for PEKK grades to remove thermal history, cooled at the rate of 5oC/min, re-heated at 10oC/min. Peak melting temperatures (T m) were obtained from the second heating cycle.Rheological AnalysisBased on the lower and the upper limit of temperatures determined by DSC and TGA analysis, three and four processing temperatures were chosen with 15oC intervals for PEKK and PPS grades respectively for rheology tests.

All rheological tests were done on TA Instruments DHR-2 instrument fitted with 25mm parallel plate geometry. The dried pellets were directly melted on the plates using melt rings. Oscillatory amplitude tests were first performed to identify the linear viscoelastic region for each of these materials. A suitable strain907value in the linear region (typically 1% or lower) was selected for frequency sweep tests between 0.1 and 628 rad/s. The purge gas used throughout was air.Filament Making Process and FFF ExtrusionFilaments for the FFF process were extruded using a Filabot EX2 (as shown in Figure 1) fitted with a single screw extruder. Using the temperature range identified for rheological analysis and by varying the screw speed and nozzle diameter, filaments were made from the dried pellets with a target diameter of 1.75mm. The extruded filaments were gently guided and rolled up under ambient conditions with minimum stretching.Figure 1.

The extruded filaments were gently guided and rolled up under ambient conditions with minimum stretching.Figure 1. Filament making processFor extruding the filaments on a fused filament fabrication system, a Solidoodle 3 was customized with a E3D V6 all metal hot end to reach temperatures as high as 400oC. The nozzle diameter was 0.6 mm for the neat filaments and 1 mm for the carbon fiber reinforced filaments. The print bed temperature was set to 100oC.Results and DiscussionThermal AnalysisFigure 2a and 2b show the TGA and DSC (second heating cycle) thermograms for PPS and PEKK grades, respectively. The TGA data indicate that all three grades of PPS are stable at least up to 400oC (less than 2% weight loss) and all PEKK grades are stable at least up to 500oC (less than 2% weight loss). This sets the upper processing temperature limit.

This sets the upper processing temperature limit. For the lower processing temperature limit, peak melting temperatures were obtained in the range of 280oC to 285oC for PPS grades and 359oC to 362oC for PEKK grades.From these limits, four processing temperatures were selected above the melting point for the PPS grades, specifically: 300oC, 315oC, 330oC and 345oC. Likewise, three candidate processing temperatures above the melting point were identified for the PEKK grades, specifically: 375oC, 390oC and 405oC.908(a)(b)(c)(d)Figure 2. TGA analysis of (a) PPS grades and (b) PEKK grades and DSC thermograms showing the second heating cycle for (c) PPS grades and (d) PEKK gradesRheological AnalysisFigure 3a and 3b show the variation of complex viscosity with frequency, temperature, and filler content for the various PPS and PEKK grades, respectively.

For PPS grades (Figure 3a), addition of fillers increase viscosity at all frequencies, with the magnitude of increase being greater for higher filler loadings. For viscosity variation with frequency, at all the chosen temperatures, the viscosity of neat PPS shows only a small dependence on frequency. However, for the filled grades, shear thinning of the material is much stronger. For temperature dependence of viscosity, for neat PPS, viscosity drops with change in temperature from 300oC to 315oC, but beyond that, an increase in the temperature does not change viscosity significantly for the entire frequency range tested. For PPS with 40 wt. carbon fiber (CF), temperature dependence of viscosity is very low for the entire range, and there is no significant difference for 50 wt.% CF at higher frequencies. However at lower frequencies, viscosity increases with an increase in temperature.

However at lower frequencies, viscosity increases with an increase in temperature. One of the possible reasons for this could be the occurrence of structural changes in the material due to chain scission and cross-linking reactions [10] at high temperatures. The magnitude of increase depends on the degree of crosslinking and shear rate. Crosslinking reactions increase viscosity. However, the newly formed structure could be sensitive to shear, thereby lowering the viscosity with increasing shear rate.909In the case of PEKK (Figure 3b), the addition of fillers increases viscosity for the entire frequency range studied (similar to PPS grades), with viscosity for 40 wt.% CF PEKK being greater than 30 wt. % CF PEKK. Here, all the three grades exhibit shear thinning behavior and the extent of shear thinning is greater for the filled grades. The temperature dependence of viscosity is very low for neat PEKK and for PEKK with 30 wt.% CF.

The temperature dependence of viscosity is very low for neat PEKK and for PEKK with 30 wt.% CF. In the case of PEKK with 40 wt.% CF, viscosities at 375oC and 390oC show a very similar trend. However, for 405oC, at low frequencies, viscosity increases by about an order of magnitude. This could be due to some structural changes induced in the material at temperatures above 400oC.(a)(b)Figure 3. Frequency sweep tests for the different grades of (a) PPS and (b) PEKK at the candidate processing temperaturesThe frequency sweep data indicate that for filled grades of both PPS and PEKK, shear rate is a much more effective parameter for controlling viscosity than temperature. Such knowledge of the effect of process parameters and fillers on flow properties can be useful while processing such new materials on the FFF system.Filament Making Process and FFF ExtrusionFilament extrusion was successful for all three chosen grades of PPS and PEKK as shown in Figure 4.

Table 1 indicates the extrusion temperature used for the Filabot and the typical diameter of the filaments.910Figure 4. Filaments of (a) Neat PPS, (b) 40 wt.% CF PPS, (c) 50 wt.% CF PPS, (d) Neat PEKK, (e) 30 wt.% CF PEKK, (f) 40 wt.% CF PEKKTable 1. Filament extrusion temperatures and diameterMaterial Neat PPS PPS 40 wt.% CF PPS 50 wt.% CF Neat PEKK PEKK 30 wt.% CF PEKK 40 wt.% CFExtrusion temperature (oC) Filament diameter (mm)300 345 345 370 375 3751.5-1.9 1.4-1.65 1.4-1.65 1.5-1.75 1.6-1.9 1.5-1.9The extruded filaments were deposited on the customized Solidoodle 3 to determine if these materials could be extruded through an FFF nozzle and deposit a few layers as shown in Figures 5a and 5b. This study demonstrated the deposition of neat PPS and PEKK and also the fiber reinforced grades of PEKK. The fiber filled grades of PPS were not successfully deposited due to filament feed issues (rough and rigid filaments), not rheological limitations.

The fiber filled grades of PPS were not successfully deposited due to filament feed issues (rough and rigid filaments), not rheological limitations. Table 2 indicates the processing conditions for FFF extrusion and deposition.(a) (b)Figure 5. (a) FFF extrusion of neat PPS and (b) Deposited layers of neat PEKK911Table 2. Extrusion temperature on the FFF systemMaterialFFF ExtrusionNeat PPS280-290oCNeat PEKK390-400oCPPS 40 wt.% CFFilament feed issuesPPS 50 wt.% CFFilament feed issuesPEKK 30 wt. % CF390- 400oCPEKK 40. wt% CF390- 400oCTypically, for most FFF or FDM processes, shear rates in the nozzle range between 100-200 s-1 [3,11].

% CF390- 400oCPEKK 40. wt% CF390- 400oCTypically, for most FFF or FDM processes, shear rates in the nozzle range between 100-200 s-1 [3,11]. To relate complex viscosities obtained from the frequency sweep tests to viscosities at FFF shear rates, complex viscosities for each of the grades of PPS and PEKK in the frequency range of 100-250 rad/s were identified using the Cox-Merz rule (equation 1)for(1)∗(𝜔𝜔) where represents frequency [12]. Using this relation, the complex viscosities and frequencies can be 𝜔𝜔 related to steady shear viscosities at FDM shear rates.𝜂𝜂(𝛾𝛾̇) = 𝜂𝜂𝛾𝛾̇ = 𝜔𝜔 is complex viscosity,is steady shear viscosity,represents shear rate, and∗𝜂𝜂𝜂𝜂𝛾𝛾̇To correlate the FFF extrusion conditions of PPS and PEKK with the viscosity data from rheological studies, the viscosities of these materials were compared with that of acrylonitrile butadiene styrene (ABS), a successfully printed FDM material, in the frequency range of 100-250 rad/s.

For ABS, frequency sweep tests were performed using the filaments (neat and 40 wt.% CF) used previously to deposit these materials on a Solidoodle 3 at 205oC in a study by Tekinalp et.al. [13].Figures 6a and 6b represent frequency sweep data for neat materials (PPS and PEKK) and carbon fiber reinforced grades of PEKK respectively at different temperatures, indicating the viscosity range for these materials at the typical FFF shear rates (here, frequency) along with the complex viscosity range for neat ABS and ABS with 40 wt.% CF in the same frequency range (measured at 205oC – the FFF deposition temperature).912(a)(b)Figure 6. Frequency sweep comparison for (a) neat PPS and PEKK with neat ABS and (b) CF filled PEKK with CF filled ABSIn comparison with neat ABS, between 100-250 s-1, the viscosity of neat PPS and neat PEKK is lower. This indicates that neat PPS and PEKK can flow through a similar FFF system as that of ABS without clogging at the mentioned temperatures.

This indicates that neat PPS and PEKK can flow through a similar FFF system as that of ABS without clogging at the mentioned temperatures. For fiber reinforced PEKK systems, the viscosity of 30 wt.% CF PEKK is lower than ABS 40 wt. % CF. For 40 wt.% CF PEKK, at the extrusion temperature, the viscosity is in the upper limit of that of filled ABS, indicating that these materials with similar viscosities can extrude through the FFF nozzle.It should also be noted that the comparisons made here involve some approximationsand the following parameters are to be considered as well:1. The applicability of Cox-Merz rule to the fiber filled systems used here is an approximation. Previous studies [14] on LLDPE filled with glass fiber have shown that the dynamic viscosity data for fiber reinforced systems is higher than steady shear viscosity obtained by capillary rheometry.

Previous studies [14] on LLDPE filled with glass fiber have shown that the dynamic viscosity data for fiber reinforced systems is higher than steady shear viscosity obtained by capillary rheometry. This was attributed to the reduction in steady shear viscosity due to fiber orientation along the flow direction in a capillary rheometer, which does not happen in the case of oscillatory frequency tests done at low strains.2. The viscosities compared are at the typical steady state FDM shear rates. It is assumed the material would be able to flow with viscosities in this range, provided the torque of the extrusion system is high enough to get the material flowing initially (overcome zero shear viscosity which can be much higher than the viscosities shown in the frequency sweep tests). Conclusions and Future WorkThis study was the very first attempt at understanding rheological properties of the chosen grades of PPS and PEKK and correlating it to FFF processing conditions.

Conclusions and Future WorkThis study was the very first attempt at understanding rheological properties of the chosen grades of PPS and PEKK and correlating it to FFF processing conditions. This work has demonstrated that rheological analysis can be helpful in identifying appropriate FFF processing conditions and understanding the key control parameters. The effect of factors such as temperature, shear rate, and fillers on viscosity have been studied for both neat and carbon fiber reinforced PPS and PEKK grades. Filaments for all the chosen materials were successfully extruded from pellets, and four grades were deposited on a customized desktop913sized FFF system.

Filaments for all the chosen materials were successfully extruded from pellets, and four grades were deposited on a customized desktop913sized FFF system. Also, the rheological properties of these candidate materials were compared with that of a common FDM material (ABS).Future work will involve comparisons of this data with steady shear tests using a capillary rheometer, understanding the effect of filler size on flow, improvements to the filament making process, and correlating experimental data with theoretical models.AcknowledgmentsResearch sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05- 00OR22725 with UT-Battelle, LLC. Materials for this work were provided by Arkema Inc. and Techmer ES. The authors also gratefully acknowledge Dr. Mark Aubart, Timothy Spahr and Dr. David Liu from Arkema Inc.

Materials for this work were provided by Arkema Inc. and Techmer ES. The authors also gratefully acknowledge Dr. Mark Aubart, Timothy Spahr and Dr. David Liu from Arkema Inc. and Alan Franc from Techmer ES for their support and guidance with this research.References1. Special Chem. Retrieved from http://omnexus.specialchem.com/selection-guide/high- temperature-thermoplastics.2. Spruiell, J. E. (2005). A review of the measurement and development of crystallinity and its relation to properties in neat poly (phenylene sulfide) and its fiber reinforced composites (No. ORNL/TM-2004/304). ORNL.3. N. Turner, B., Strong, R., & A. Gold, S. (2014). A review of melt extrusion additive I. Process design and modeling. Rapid Prototyping manufacturing processes: Journal, 20(3), 192-204.4. Jarrousse, G. (2004).

Wasley1,*,†, J. Li1,* , D. Ta2, J. Shephard2, J. Stringer3, P. Smith3, E. Esenturk4, C. Connaughton4, and R. Kay51The Wolfson School of Mechanical, Electrical and Manufacturing Engineering, Loughborough University, Loughborough, UK 2Institute for Photonic & Quantum Sciences, Heriot-Watt University, Edinburgh, UK 3Department of Mechanical Engineering, University of Sheffield, Sheffield, UK 4Warwick Mathematics Institute, University of Warwick, Coventry, UK 5Institute of Medical and Biological Engineering, School of Mechanical Engineering, University of Leeds, Leeds, UK *T.J.Wasley@lboro.ac.ukAbstractAdditive Manufacturing (AM) processes can facilitate the rapid iterative product development of electronic devices by optimising their design and functionality. This has been achieved by combining two additive manufacturing processes with conventional surface mount assembly to generate high resolution embedded multilayer electronic circuits contained within a 3D printed polymer part.

Bottom-up DLP Stereolithography and material dispensing of isotropic conductive adhesives have been interleaved to deposit microscale features on photopolymer substrates. The material dispensing process has demonstrated the high density deposition of conductors attaining track widths of 134µm and produced interconnects suitable for directly attaching bare silicon die straight to the substrate. Interconnects down to a diameter of 149µm at a pitch of 457µm have been realized. In addition, this research developed a novel method for producing high aspect ratio z-axis connections. These were simultaneously printed with the circuit and component interconnects by depositing through-layer pillars with a maximum aspect ratio of 3.81.

These were simultaneously printed with the circuit and component interconnects by depositing through-layer pillars with a maximum aspect ratio of 3.81. Finally, a method to accurately embed the packaged circuit layer within the printed part has been employed using bottom-up stereolithography.Keywords: additive manufacturing; process integration; multilayer embedded electronics; DLP stereolithography; material dispensing; flip chip packagingIntroductionAdditive manufacturing technologies have been proven to provide substantial advantages over traditional manufacturing methods including greater process versatility, achievable geometric complexity and digitally driven fabrication capability. The application of theseThese authors contributed equally to this work.

The application of theseThese authors contributed equally to this work. † Author to whom any correspondence should be addressed.1838processes in electronics production helps to overcome the template driven nature of conventional electronics manufacturing processes that result in a lack of customisation and long pre- production timescales. Individual AM processes and materials are typically limited in their processing capabilities. Electronic devices that require the multi-material compatibility and components can therefore be printed by interleaving two separate processes capable of depositing both dielectric and conductive materials.

Such a process capability would therefore provide a larger range of applications for embedded sensor systems, microfluidics and bespoke electronic devices.Successful attempts to hybridise AM processes have been previously been conducted for the fabrication of electronics including the production of substrates by Stereolithography (SL), Ultrasonic Consolidation (UC) [1], Fused Deposition Modelling (FDM) [2], material jetting [3] and Selective Laser Sintering (SLS) [4] and; the deposition of conductors by aerosol jetting, ink jetting and dispensing based direct writing techniques [5].functional components using shape deposition manufacturing for build 3D end-use products was first explored in 1992 [6] and further adopted by Kataria and Rosen [7] who proposed a technique for embedding functional inserts in SL parts.

The University of Texas at El Paso introduced a hybrid AM setup combining material dispensing of low viscosity conductive inks with laser-based top-down stereolithography [8][9][10]. Photopolymer substrates were produced with surface channels and trenches into which conductive ink could be deposited and laser cured and, components positioned and individually embedded. In addition, through holes were created during the SL process which were subsequently pumped with conductive ink [11]. These cavities inhibit the placement of small surface mount devices and therefore this technique does not lend itself to the production of high density multilayer electronics. FDM was then applied to this hybrid concept to build a 2.5D circuit board [12]. Niese et al.

FDM was then applied to this hybrid concept to build a 2.5D circuit board [12]. Niese et al. [13] presented a concept for encapsulation of PCBs and fabrication of metallic channels to route a power source using an Additive 3D Molded Interconnect Device (ADDMID) process, with Stereolithography using a doped photopolymer replacing injection moulding to create polymer parts for laser processing.Thefeasibility of embeddingIn 2016, Voxel8 released a material extrusion system with multi-material capability coupling Fused Filament Fabrication (FFF) and dispensing within a single printer [14]. This allows a polymer substrate and conductive material to be sequentially deposited creating a polymer package with internal electronics. Stratasys and Optomec also demonstrated direct writing onto conformal FDM substrates by aerosol jetting, printing circuits, sensors and antennas onto an unmanned aerial vehicle (UAV) [15].

Stratasys and Optomec also demonstrated direct writing onto conformal FDM substrates by aerosol jetting, printing circuits, sensors and antennas onto an unmanned aerial vehicle (UAV) [15]. FFF and FDM do however demonstrate disadvantages such as low resolution, weak adhesive strength between layers, poor surface finish due to stair stepping and difficulty producing hermetic structures [16].Finally, non-polymer based technologies have been investigated for use in hybrid additive manufacturing of electronics. Ultrasonic consolidation has been used to embed1839electronics within three dimensional metal structures [17][18][19] and, ink jetting has been used to print conductive inks into additively manufactured ceramic substrates [20].This paper introduces a new digitally driven method developed for the fabrication of functional multilayer electronics devices.

Conductors can be printed directly onto the surface of a photopolymer substrate, with no need for features to control the deposition of material or position of components. Dispensed circuitry and low temperature assembled and packaged electronic components can be embedded within a thick layer of photopolymer.

Dispensed circuitry and low temperature assembled and packaged electronic components can be embedded within a thick layer of photopolymer. During this work investigations have been conducted to:characterise the dispensing process for the production of conductive traces, pillars and interconnectsoptimisation of the electrical properties of the printed conductors • analysis of the interaction between conductive materials and photopolymers • proving the application of this manufacturing technique for flip chip packaging Finally demonstrators in the form of multilayer flashing 555 timer circuits are presentedto prove the technological capability.Hybrid Additive Manufacturing ProcessFigure 1(a)-(h) shows the hybrid manufacturing process flow beginning with the fabrication of acrylic-based photopolymer substrates using bottom-up DLP stereolithography.

A mid-process developing technique including ultrasonic agitation is then utilised to decontaminate and optimise the substrate surface quality for dispensing. This process is compatible with surface mount assembly, a consequence of the utilization of isotropic conductive adhesives (ICAs) to form conductive traces, novel freestanding through-layer pillars and component interconnects allowing surface mount devices (SMDs) to be deposited directly onto the bond pads. After surface mount assembly, the ICA is thermally cured, encapsulated within a thick layer of photopolymer and a new cavity fabricated to locate the next layer of circuitry.1840Figure 1.Hybrid AM Process ChainMultilayer Photopolymer Substrate FabricationThe Stereolithography system utilises a DLP pico-projection mechanism with a Digital Micromirror Device (DMD) [21] dynamic mask and an ultraviolet (UV) 405nm LED source.

This light source is arranged in a bottom-up orientation, with the projector focused on the base of the material vat and an average light intensity of 0.962mW/cm2, requiring a 10 second exposure to cure a 100µm thick layer of photopolymer. DLP control software also facilitates an increase in the light intensity to 3.16mW/cm2 however; this consequently causes faster deterioration of the anti-stiction PDMS layer on the vat base and has been proven to degrade the reflective coating of the DMD. The setup provides a resolution of 109µm over a 140mm x 87mm projection area and, customisable z-axis layer thickness. This approach results in a faster build time than vector scanning methods and, when combined with the bottom-up projection orientation results in a substrate surface roughness (Ra) of 2µm and an average change in surface height of 6µm over a 15mm x 15mm sample area. This high quality is achieved due to its proximity to the PDMS layer on the base of the vat during curing.

This high quality is achieved due to its proximity to the PDMS layer on the base of the vat during curing. The photopolymer is an acrylic-based broad-band resin sensitive to wavelengths of light up to 440nm and has a glass transition temperature of ~100°C.1841Mid-Process Substrate DevelopingTo facilitate the integration of Stereolithography with additional AM processes, decontamination of the substrate surface is required to ensure the successful deposition of conductive materials onto the photopolymer. Ultrasonic agitation in a solvent was identified as the most efficient method of thinning and removing excess liquid resin from the substrate surface. Specimens fabricated using the broad-band acrylic-based photopolymer exhibited surface cracking after exposure to a number of common solvents compromising their surface quality due to incomplete polymer crosslinking.

Alternative solvents and additional photoresist processing steps including relaxation and thermal baking in addition to an UV oven exposure were investigated for their effect on the substrate surface quality [22].Photopolymer samples were exposed for 5 minute, 30 minute, 2 hour and 24 hour timeperiods to eight different solvents:1. Isopropyl Alcohol (IPA) 2. Methanol 3. Acetone 4. D-Limonene 5. Xylene 6. Ethanol 7. Toluene 8. TetrahydrofuranD-Limonene and IPA showed no degradation of the sample or its features however, surface cracking was demonstrated on every specimen.

D-Limonene 5. Xylene 6. Ethanol 7. Toluene 8. TetrahydrofuranD-Limonene and IPA showed no degradation of the sample or its features however, surface cracking was demonstrated on every specimen. Relaxation and thermal baking resulted in a small reduction in the frequency of cracking however, 30 minutes flood exposure to light in a UV oven after solvent rinsing in D-Limonene resulted in elimination of the cracking as shown in Figure 2.Figure 2.Comparison between unprocessed and processed substrate surfaces1842This is a result of the continuation of crosslinking and therefore prevention of solvent penetration in the substrate. Optimal surface quality was therefore achieved by 2 minutes of ultrasonic agitation in D-Limonene and 30 minutes flood exposure in a UV oven.Dispensing ConductorsA Musashi Shotmaster 500 dispensing system with a Musashi SuperΣ®CMII digital pressure controller was used to print conductors on photopolymer substrates.

The system is comprised of five main parts:a 3-axis CNC router • a computer controller • a digitally controlled dispenser • an alignment camera • a laser positioning feedback systemWhen combined, these parts are capable of depositing high viscosity, heavily loaded ICAs with viscosity up to 1,000Pa.s with printing pressures between 30kPa to 500kPa and speeds of 0.1mm/s to 300mm/s at a movement resolution of ±5µm. Printing is controlled in 3 axes and a laser is used to map the surface of the substrate, automatically adjusting the nozzle height to maintain a constant print gap.Two commercially available silver based isotropic conductive adhesives, Epotek E4110- PFC and Epotek EJ2189 with maximum particle sizes of ≤20µm and ≤45µm [23][24], were selected due to their rheological properties and low temperature curing requirements, making them compatible with SL photopolymer materials.

Control over the deposition of these ICAs can be achieved through a combination of five parameters: 1) viscosity of the dispensed material; 2) inner diameter (ID) of the dispensing nozzle; 3) extrusion pressure; 4) print speed and; 5) print gap (distance between the nozzle tip and substrate).Conductors have been identified for fabrication using the combination of material dispensing and ICAs. Conductive traces were printed to connect multiple discreet components and integrated circuits (ICs) and, bond pads deposited on which these components can be placed. In addition, to facilitate the production of multilayer embedded electronics, a new method for the production of through layer connections has been developed.

In addition, to facilitate the production of multilayer embedded electronics, a new method for the production of through layer connections has been developed. Freestanding pillar structures were printed in the z-axis in sequentially deposited layers, to produce features capable of spanning the maximum 2mm distance between two adjacent layers of circuitry.Electronic PackagingSurface mount assembly (SMA) is highly advantageous as it is the most commonly used packaging method due to its low cost, high throughput and high density capability. This technique allows components to be attached directly onto surface of printed circuit boards, functionality that is also demonstrated by this hybrid process as components can be mounted1843onto dispensed ICA interconnects on the photopolymer surface.

The achievable printing resolution of the dispensing process determines the SMD size that can be mounted onto the deposited bond pads, with this process demonstrating the packaging of both 1206 and 0603 SMDs. In addition, ICs with pins at pitches of 1.27mm and 500µm have been mounted. The height of these components determines the necessary embedding layer thickness identified as 2mm and 1.2mm for larger and smaller devices respectively and therefore, the height of the conductive pillars.The electronics industry is driven by demand for the minimisation of circuitry and therefore, this hybrid AM process chain has been adapted, as shown in Figure 2, for use in flip chip packaging. This packaging technique is an evolution of SMA where bare microchips, predominantly semiconductors, are attached face down onto a substrate reducing weight, making the device thinner and achieving a higher input/output density due to smaller pitch between an array of bond pads [25].

This functionality has been proven using daisy chain patterned bare die, requiring alternate bond pads to be contacted to create a complete electrical connection around its perimeter. Substrates were fabricated with 600µm x 200µm x 200µm (Length x Width x Depth) channels in the surface corresponding to the daisy chain pattern, ultrasonically agitated in solvent and filled with an ICA before low temperature thermal curing. The surface was then lapped to planarise and remove any excess conductive materials from the substrate surface. ICA interconnect bumps were then deposited on either end of each filled channel before the bond pads on the bare die were aligned with the bumps, the bare die placed and a bonding force applied to the chip.Figure 3.Flip chip packaging via hybrid AMThe substrate fabrication technique allows embedding of the electronics in a thick layer of photopolymer.

This provides a high quality surface with exposed through layer connections for the dispensing of subsequent layers of circuitry. One benefit of the Stereolithography process is its ability to produce geometrically complex structures, an advantage that can be maintained after hybridisation by fabricating external sidewalls surrounding a cavity in which circuitry can be deposited as shown in Figure 3. These cavities can then be submerged in liquid photopolymer facing upwards causing de-gassing and SMD underfilling. The substrate is inverted and immersed in the material vat, pressing the tips of the conductive pillars into the PDMS layer on1844its base.

The substrate is inverted and immersed in the material vat, pressing the tips of the conductive pillars into the PDMS layer on1844its base. Finally, the filled cavity is selectively exposed to UV light for a prolonged period of time after which a subsequent cavity is fabricated for the next layer of circuitry.Figure 4.Process of embedding electronics using StereolithographyCharacteristics of Dispensed ConductorsPrinting parameters can cause significant differences in the size and quality of the conductors and have been characterised in their production of conductive traces, z-axis through layer pillars and flip chip interconnects.

The objective was to print repeatable features of the smallest possible size enabling the size of the packaging and circuit footprint to be minimised.Conductive TracesRegardless of the resistivity of the conductive adhesive, line resistance can be controlled through the track geometry, with increased cross sectional area reducing the line resistance. Tracks were printed on the substrate surface in both ICA materials to determine the effect of printing pressure, print speed and nozzle size on their behaviour and appearance. A 10x objective lens was used to obtain profile data of sets of five conductive traces printed in different combinations of parameters. The width and peak height of these tracks were then extracted from this data. Higher viscosity E4110-PFC conductive adhesive produced a smoother surface finish and more semi-circular profile.

The width and peak height of these tracks were then extracted from this data. Higher viscosity E4110-PFC conductive adhesive produced a smoother surface finish and more semi-circular profile. Data presented in Table 1 shows that in higher volume deposits E4110-PFC showed a higher aspect ratio than lower viscosity EJ2189 due to material slumping however, as the nozzle inner diameter is decreased and the dispensed volume reduces, this slumping effect is reduced resulting in higher aspect ratios printed in EJ2189.1845Table 1. Comparison of aspect ratios produced by E4110-PFC and EJ2189 ICAsNozzle Size (µm)100 200 250E4110-PFCAvg. Width (µm) 234 370 419Aspect Ratio 0.278 0.315 0.387EJ2189Avg. Width (µm) 218 363 420Aspect Ratio 0.300 0.306 0.356The smallest tracks produced were 134µm in width and 38µm in height. These were printed in the lower viscosity EJ2189 ICA at a pressure of 140kPa, 7mm/s print speed and 100µm nozzle ID.

These were printed in the lower viscosity EJ2189 ICA at a pressure of 140kPa, 7mm/s print speed and 100µm nozzle ID. The lowest trace dimensions in E4110-PFC through a 100µm nozzle ID were 170µm width and 42µm height printed at 300kPa and 7mm/s print speed. These track widths enable fine pitch conductors to be printed reducing the size of compatible electronic components and the corresponding package.The lowest resistivity possible is desired to reduce the line resistance of conductors while minimising their size. Conductive adhesives achieve their conductivity through two mechanisms, shrinkage of the resin part of the epoxy formulation increasing contact between particles or flakes of the conductive filler materials and, evaporation of lubricants further reducing the distance between conductive particles [26]. Three different low temperature thermal curing regimes listed in Table 2 had their effect on material resistivity measured.

Three different low temperature thermal curing regimes listed in Table 2 had their effect on material resistivity measured. The glass transition temperature of the photopolymer limited the processing temperature to 100°C.Table 2. Manufacturer’s specified and chosen experimental curing regimes for ICA materialsExperimental Regimes EJ2189 Spec Temp (°C) Time (hrs) Temp (°C) Time (hrs) Temp. (°C) Time (hrs)E4110-PFC SpecLow Med High45 80 1206 3 123 80 10072 3 145 80 100Optical microscopy was used to capture the topology of track samples produce in both ICAs and cured using all three regimes. The resulting clouds of data points were processed using a custom MATLAB® code designed to compute the average cross sectional area of each measured length of track.

The resulting clouds of data points were processed using a custom MATLAB® code designed to compute the average cross sectional area of each measured length of track. Track resistance (R), track length (L) and cross sectional area (A) were then substituted into equation 1 to calculate the resistivity (ρ):(1)𝐿𝐿𝑅𝑅 = 𝜌𝜌𝐴𝐴 In both materials, higher temperature processing at 100°C for 1 hour resulted in lower resistivity with EJ2189 demonstrating a resistivity of ~2.8x10-4 Ω.cm compared to ~4x10-4 Ω.cm in E4110-PFC. This difference between the two ICAs is shown across all curing regimes in18466 3 1Figure 4.

This difference between the two ICAs is shown across all curing regimes in18466 3 1Figure 4. One explanation for this lower resistivity in EJ2189 is its use of flakes as opposed to particles, creating a greater contact area between adjacent flakes.Figure 5.Effect of different curing regimes and embedding on ICA volume resistivityAfter embedding in a thick layer of photopolymer, both ICAs exhibited an increase in resistivity however, the average increase across all regimes in EJ2189 was just 8% compared to 23% in E4110-PFC. After curing at the highest temperature some degradation of the SL substrate was evident and therefore, 80°C for 3 hours was adopted as the default curing regime.Conductive PillarsVertical conductive structures were introduced to realise conductive connections between the multiple layers of stacked circuitry.

The printing of these structures was facilitated by the viscosity of the ICA materials and the three axis capability of the dispensing equipment. Figure 5 shows the clockwise square helix design of the conductive pillar and proportional alterations made to the pillar dimensions when printed through the three available nozzle IDs.1847Figure 6. Conductive pillar design and dimensionsMaterial slumping demonstrated during track deposition was also exhibited during pillar printing and, due to the multiple layers of successive dispensing and additional weight of these layers, the effect was exaggerated. Higher viscosity E4110-PFC slumped less and therefore produced higher aspect ratio, taller and narrower structures. Printing characterisation was conducted using printing pressure, print speed and nozzle ID as variables to determine the optimal combination of parameters for each square helix design shown in Figure 5.

Once the printing behaviour and materials had been assessed, additional layers could be added to each design one by one, increasing the height until the printed structure collapsed at its height limit.Figure 6 showed that E4110-PFC created the tallest structures using each of the three nozzle sizes. 16, 18 and 24 layers were printed through 250µm, 200µm and 100µm nozzle IDs resulting in aspect ratios of 3.81, 3.23 and 3.68. The 100µm nozzle ID produced a pillar in excess of 2mm in height and 553µm wide, making it suitable to span the thickest proposed 2mm embedding layer in this investigation while minimising the footprint on the substrate.

The maximum aspect ratio achieved using lower viscosity EJ2189 was 2.41 with a height of 1.175mm over 17 layers.1848Figure 7.Topologies of maximum pillar heights produced using each nozzle IDFlip Chip InterconnectsThe deposition of flip chip interconnects onto a substrate surface is conducted by timed point deposits, differing from trace and pillar dispensing as there is no nozzle movement during the application of pressure to the ICA. For the purposes of this process characterisation, print gap and pressure actuation time were introduced as variables and, print speed removed. The lower viscosity of EJ2189 demonstrated better wetting behaviour than E4110-PFC as it separated from the nozzle more easily, allowing smaller volumes of ICA to be deposited and therefore smaller bumps to be produced. Figure 5 shows the smallest consistent interconnects produced with an average width of 149µm and height of 80µm.

Figure 5 shows the smallest consistent interconnects produced with an average width of 149µm and height of 80µm. They were printed with a combination of 0.2s pressure actuation time, 30kPa pressure and 60µm print gap in EJ2189.Figure 8.Smallest flip chip interconnects with consistent profile1849The adaptation of this hybrid AM process for flip chip packaging was validated using daisy chain connections and a bare silicon test die. This set-up provides a mechanism by which contact resistance between adjacent connections can be measured. The sourced bare die were only available pre-bumped with solder, features made unnecessary by the ICA interconnects however, the success of this flip chip packaging process could still be proven. Figure 8a shows redistribution traces designed into the substrate surface allowing the resistance of each individual connection to be measured.

Figure 8a shows redistribution traces designed into the substrate surface allowing the resistance of each individual connection to be measured. The resistance of each connection reduced from 5.2Ω to 2.3Ω by adding a 0.08N bonding force and removing the solder bumps that increased the length of the conductive pathway. The solder removal also reduced the chip standoff from 160µm (Figure 8b) to 80µm, indicating the distance between the surface of the substrate and underside of the bare die.In traditional electronics assembly, an insulating polymer is used to fill the void between a chip and the substrate protecting interconnects from stresses such as vibration or shock [27]. The embedding process introduced as part of this hybrid AM technique has also been used to underfill flip chip packaged bare die with photopolymers, with no measurable change in contact resistance.

The embedding process introduced as part of this hybrid AM technique has also been used to underfill flip chip packaged bare die with photopolymers, with no measurable change in contact resistance. Figure 8c and Figure 8d shows cross sectional micrographs of successfully underfilled bare die at a standoff height of 160µm. The same process conducted on die with an 80µm standoff showed bubbles trapped between the substrate and chip. Ultrasonic agitation is one technique that could be investigated further for used in removing this trapped air.Figure 9.Flip chip packaging and underfilling of daisy chain bare die1850DemonstratorsProcesses and methods for interleaving multiple additive manufacturing techniques have been developed. These processes have been designed to produce conductive tracks, interconnects and z-axis connections. A flashing 555 timer circuit was selected to demonstrate the application and integration of these processes.

A flashing 555 timer circuit was selected to demonstrate the application and integration of these processes. It provides a simple and visual exhibition of this hybrid manufacturing capability and, embedding of a variety of surface mount devices including discrete components (resistors and capacitors), an actuator (LED) and a small outline integrated circuit (555 timer chip).Feasibility DemonstratorsInitial demonstrators were designed to evaluate the effectiveness of the integration of multiple processes and therefore no not demonstrate the geometric capabilities of this manufacturing technique. Figure 9 shows single, double and triple layer feasibility demonstrators, all of which flashed as designed when a power source was applied, proving that ICA conductors and SMDs could be successfully embedded on multiple layers within a photopolymer substrate.

In addition, the triple layer device was the first to utilise a miniaturised circuit layout decreasing the footprint of printed conductors from 25mm x 15mm to 9mm x 9mm and halving the size of SMDs from 1206 to 0603 (imperial code, resistor dimension of 1.6mm x 0.8mm x 0.8mm) and, reducing the size of IC pin pitch from 1.27mm on a Mini Small Outline Package (MSOP) to 500µm on a Thin Shrink Small Outline Package (TSSOP).Figure 10.Single, double and triple layer feasibility demonstrators1851Pyramid DemonstratorsFollowing the demonstration of successful fabrication and function of the integrated AM processes, the capability of this technique to produce high complexity packaging geometries with internal electronics was explored. A pyramid shape was chosen as it displays a constant change in external geometry that could only be achieved using this technique.

A pyramid shape was chosen as it displays a constant change in external geometry that could only be achieved using this technique. Initially, a 30mm x 30mm pyramid base was designed (Figure 10a) and built (Figure 10b) with a cavity in which circuitry could be deposited before embedding and, the subsequent production of the next cavity enclosed within external side walls. The characterisation of conductor deposition allowed a reduction in size of the pyramid base (Figure 10c) to 15mm x 15mm with ICAs deposited using a 100µm ID nozzle.

Finally, this process cycle was repeated to completion of the circuit at which point the peak was fabricated to fully encapsulate the electronics and complete the fully functioning pyramid demonstrators (Figure 10d) that flash intermittently when connected to a power source.Figure 11.Pyramid a) CAD design, b) 30mm x 30mm base, c) 15mm x 15mm base and d) fully functionaldemonstratorConclusionsA novel hybrid manufacturing process has been presented showing the integration of Stereolithography and direct writing technologies with a mid-process developing procedure and conventional SMA of discrete electronic components, actuators and integrated circuits.A DLP Stereolithography system in a bottom-up orientation was applied to the fabrication of electronic substrates and the packaging of electronics within thick layers of cured photopolymer.

Curing the resin in contact with the base of the vat results in high quality surface finishes on both the initial base substrate and subsequent thick embedding layers.1852High viscosity, low temperature curing isotropic conductive adhesives were combined with material dispensing to produce conductors including tracks, interconnects and, novel freestanding conductive pillars designed to create through layer conductive connections. Surface mount devices could be mounted directly onto the printed bond pads before they were thermally cured at 80°C for 3 hours, resulting in a higher conductivity than specified by the manufacturer. When embedded within the polymer matrix, SMDs were underfilled, tips of conductive pillars pushed into the soft PDMS layer on the vat base and, only an 8% increase in ICA resistivity was observed.

When embedded within the polymer matrix, SMDs were underfilled, tips of conductive pillars pushed into the soft PDMS layer on the vat base and, only an 8% increase in ICA resistivity was observed. In addition to surface mount packaging, chip scale flip chip packaging has been conducted using a combination of Stereolithography and dispensing techniques to produce 600µm x 200µm x 200µm conductive trenches with 149µm flip chip interconnects. Bare die have also been successfully underfilled using the liquid photopolymer to reduce stresses on the ICA interconnects.This hybrid additive manufacturing process has been successfully used to fabricate geometrically complex pyramid demonstrators with internal triple layer flashing 555 timer circuits and, package flip chip bare die.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAdditive Manufacturing of High-Entropy Alloys – A ReviewWenyuan Cui*, Xinchang Zhang*, Frank Liou*Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO, 65409AbstractHigh-entropy alloys have attracted increasingly interest due to their unique compositions, microstructures and mechanical properties. Additive manufacturing has been recognized as a promising technology to fabricate the high-entropy alloys in the recent years. The purpose of this paper is to review the current research progress in high-entropy alloys by additive manufacturing process. It will first highlight the important theory of the high-entropy alloys. The next aspect is to summarize current additive manufacturing methods applied for the high entropy alloys.

It will first highlight the important theory of the high-entropy alloys. The next aspect is to summarize current additive manufacturing methods applied for the high entropy alloys. At last, the correlation between the microstructures and the mechanical properties of the high-entropy alloys will be examined and discussed.Keywords: Additive Manufacturing; High-Entropy Alloys; Microstructures; Mechanical PropertiesIntroductionHigh-entropy alloys (HEAs) were first proposed by Yeh et al. and have attracted increasingly attention from the researchers and scientists [1]–[10]. The conventional metallic alloys usually are choosing one principal element with small amounts of other elements added to improve the properties. By contrast, HEAs are defined as alloys that have at least five principal elements and the concentration of each element is between 5 at. % to 35 at.%.

By contrast, HEAs are defined as alloys that have at least five principal elements and the concentration of each element is between 5 at. % to 35 at.%. Yeh summarized the core effects for HEAs: high-entropy effect, sluggish diffusion, severe lattice distortion and cocktail effects [1]–[3]. The high-entropy effect implies that the high mixing entropy of the elements leads to the formation of single-phase solid solution with simple crystal structures like face-center cubic (fcc) or body-center cubic (bcc) rather than intermetallic. In HEAs, the properties can be influenced by adjusting the composition change and this is called cocktail effect. It has been a challenge for the researchers to investigate the fundamental of HEAs due to the lack of the thermodynamic database and phase diagrams. Current approaches and rules to predict the phases in HEAs are based on the empirical equations like Gibbs energy, mixing entropy and mixing enthalpy [9][11].

Current approaches and rules to predict the phases in HEAs are based on the empirical equations like Gibbs energy, mixing entropy and mixing enthalpy [9][11]. Zhang and Yang proposed the parameters of W and atomic size d as a starting point to predict the formation of solid-solution phases [6], [7]. Valence electron concentration (VEC) is another criterion to decide the bcc or fcc crystal structure in HEAs [12]. These parameters have been applied in some alloys systems.The microstructures and mechanical properties of HEAs have been reported indicating their high strength, wear resistance, ductility and corrosion resistance [5], [6], [9], [13]–[15]. Most of the work utilized casting or arc melting to produce HEAs. However, it is not easy to overcome the inherent complexity and high degree of controls by the traditional manufacturing methods [16],712[17]. Moreover, it is unlikely to design the industrially suitable route to mass produce HEAs.

Moreover, it is unlikely to design the industrially suitable route to mass produce HEAs. As a result, additive manufacturing (AM) with a higher level of local process and rapid solidification rate will be a promising way to produce HEAs as engineering materials [18], [19]. Wu et al. successfully synthesized FeCoCrAlCuNix, FeCoCrAlNiTix, FeCoCrAlNi HEAs coatings free of cracks and pores by laser surface alloying [20]–[22]. Selective laser sintering (SLM) was chosen to produce FeCoCrNi HEA and it showed high strength and ductility comparable to the stainless steel [16]. AlCoCrFeNi HEA produced by selective electron beam melting (SEBM) demonstrated six times higher fracture strength than that of a conventional engineering material [17], [23].This work aims to provide a thorough literature review on the powder feed and powder bed laser additive manufacturing process of the common HEA parts.

It will first introduce the general concept of HEAs design and unique characteristics of powder feed and powder bed laser additive manufacturing. Furthermore, the bulk of the work will focus on the evaluation of microstructures, mechanical properties together with the corrosion resistance of HEAs.Theory of High-Entropy AlloysThe prediction and the stability of the solid solutions have been a topic of significant importance in current HEAs research. Thermodynamics parameters like Gibbs energy, entropy and enthalpy among the phase competition in HEAs were delicately discussed [8], [24]. For the random solid-solution system of N component, the configuration entropy can be expressed in Eq. (1).

For the random solid-solution system of N component, the configuration entropy can be expressed in Eq. (1). Since HEAs possess more than five principal components, it increases the extent of confusion in the system and mixing entropy.A common approach is the utilization of Hume-Rothery rule, which the atomic size (d) and the mixing enthalpy (∆𝐻#$%) are the two domain factors to form substitutional solid solutions. Zhang et al. proposed the parameter d to study the relationship between the phase stability and the difference in atomic size between the N-element compositions Eqs. (2), (3) [5]–[7].∆𝑆#$% = −𝑅+ $,-𝐶$𝑙𝑛𝐶$ (1)∆𝐻#$% = 4+ $,-,$31#$%𝐶$𝐶1∆𝐻$1(2)𝛿 = 100+ $,-𝐶7(1 −𝑟$+ $,-𝑐$𝑟$)<(3)where ∆𝐻#$% are the mixing enthalpy between the A element and B element, 𝑐$ is the molar concentration of the ith component and	𝑟$is the atomic radius of the ith component. It indicates that the solid-solution formation requires 𝛿 ≤ 6.6%.Yang et al.

It indicates that the solid-solution formation requires 𝛿 ≤ 6.6%.Yang et al. established another parameter W which defines the combined effects between∆𝑆#$% and ∆𝐻#$% used to predict the solid-solution phase formation in HEA [7], [9].Ω =CD∆EDFG |∆IDFG|(4)𝑇# =+ $,-𝑐$(𝑇#)$(5)713where 𝑇# is the average melting temperatures of the N-components in the alloy system. By analyzing the equations above, Yang pointed out that W=1 is a critical value to form the solid solution. For W>1, the contribution of ∆𝐻#$% is less than that of 𝑇#∆𝑆#$% and the multi- component system consists of the solid solutions.Guo et al. studied the stability of face-center cubic(fcc) and body-center cubic(bcc) solid solutions are determined by valence electron concentration (VEC) [12]. A higher valence electron concentration (VEC³8) favors the fcc solid solutions and bcc solid solution is stabilized at a lower value (VEC<6.87). The VEC values of HEAs are expressed in Eq.

A higher valence electron concentration (VEC³8) favors the fcc solid solutions and bcc solid solution is stabilized at a lower value (VEC<6.87). The VEC values of HEAs are expressed in Eq. (6).𝑉𝐸𝐶 =+ $,-𝐶$(𝑉𝐸𝐶)$(6)where 𝑐$ is the molar concentration for the ith component, (𝑉𝐸𝐶)$ is the valence electron concentration for the individual ith composition.Table 1 illustrates the alloy composition, the calculated values of d, VEC, W and crystal structures reported by additive manufacturing. It can be seen that the formation of fcc or bcc solid solution is in good agreement with the calculated parameters for most of the cases. However, the crystal structure in a few alloys does not coincide with the parameter d. For example, FeCoCrAlNiTi1.5 and FeCoCrAlNiTi2.0, the metastable phase Ti2Ni was reported along with fcc and bcc structures.

For example, FeCoCrAlNiTi1.5 and FeCoCrAlNiTi2.0, the metastable phase Ti2Ni was reported along with fcc and bcc structures. This phenomenon could be mainly attributed to the improved solubility and a uniform compositional distribution in the rapid solidified laser coating.

Most commonly used methods in HEAs are powder feed and powder bed fusion methods. A set up of direct metal deposition (DMD) with coaxial powder feeder is shown in Figure 1. During the process, a concentrated laser beam melts the synchronously fed powders and a thin layer of the substrate. The metallic powders undergo a rapid melting and cooling process. Meanwhile, a metallurgical bonding can be formed between the substrate and the powders. AlxFeCoCrNix (x=0.3,1) HEAs were fabricated by DMD [27]. Similar method of laser cladding is also applied in FeCoCrAlCu [21], AlxFeCoNiCuCr [28]. Kunce used the direct laser-engineered net shaping (LENS) for HEA as a hydrogen storage material [29].Figure 1 Schematic diagram of direct metal deposition (DMD) process setup[30]For the laser-based additive manufacturing methods, the precursor metallic powders are melted by high energy laser or electron beam in a layer by layer manner[16], [31], [32].

A schematic selective laser melting (SLM) layout is present in Figure 2. The geometrically complex components with high dimension precision can be obtained by the SLM process without subsequent process. The inert Argon or Nitrogen is usually used as the atmosphere environment. Fujieda first reported the feasibility of this method used for AlCoCrFeNi HEA, which facilitated a high level of local process control and generated rapid solidification rates [17]. It showed the ductility and fracture strength were remarkably improved compared to the conventional methods.

It showed the ductility and fracture strength were remarkably improved compared to the conventional methods. It was further explained the preheating process led to low residual stress and significantly affected the microstructures and mechanical properties[23], [33], [34].715Figure 2 A typical selective laser melting (SLM) machine layout [31]Microstructure CharacterizationMicrostructures and mechanical properties of HEAs have been frequently investigated.As the metallic powders characteristics have critical an impact on the mechanical properties, some amount of studies have been conducted in this area. The pure elemental powders of Fe, Co, Ni, Cr and Cu (99.7% wt.% minimum) were mechanically mixed together and the particle sizes were about 50-120 µm [31] Additives of Si and Mo were added into the FeCoNiCrCu powders as master alloys in order to avoid the different melting points and densities of nonmetallic elements.

The morphology of the AlCoCrFeNi powders used for the HEA is in Figure 3. It was observed that the powder particles were spherical or nearly spherical with relatively smooth surface. But a few aggregates of the powder particles were formed due to the turbulent flow in the atomization chamber and the collision of the solidifying droplets [29].Figure 3 Morphology (a), SEM/BSE micrographs (b) of the AlCoCrFeNi powder particles[29]Figure 4 shows the SEM/BSE images of the AlCoCrFeNi HEAs cast and SEBM specimens. The images of the SEBM specimens were obtained at the cross-section perpendicular to the build direction. It is seen the cast specimens have much coarser grain sizes than the SEBM716counterparts (Figure 4 a-c).

It is seen the cast specimens have much coarser grain sizes than the SEBM716counterparts (Figure 4 a-c). The SEBM specimens present the equiaxed grains, but the more coarsening of the modulated structure was seen at the bottom of the part.Figure 4 SEM images of the (a) cast, the SEBM (b) top and (c) bottom of AlCoCrFeNi HEA [23]Mechanical PropertiesMechanical properties of HEAs have been studied from the aspects of hardness, tensile and yield strength testing. The studies indicate that the mechanical properties of additive manufactured HEAs are comparable or better than the conventional alloys [24], [35]–[37]. Moreover, the process parameters, alloying element and post-treatment may cause remarkable effect on the mechanical properties.Vickers hardness testing has been widely used in mechanical property characterization of HEAs. Vickers hardness of the most common HEAs are compared with the conventional Ni, Al and Fe and Ti based alloys in Figure 5.

Vickers hardness of the most common HEAs are compared with the conventional Ni, Al and Fe and Ti based alloys in Figure 5. The hardness varies in one alloy system. For instance, the AlxFeCoNiCuCr system, the hardness varies from 390HV to 687HV depending on the contents of Al. The hardness of several types of HEAs like AlCoCrFeCu, AlCrFeCoCuTiNi and AlCoCrFeNiTi are generally higher than the conventional alloys.717Figure 5 Hardness of most common HEAs compared with the conventional alloys using the data from [9], [20], [22], [29], [38]–[44]The selection of alloys system is critical in determining the hardness of HEAs. The hardness has the trend of increase with the Ni content in Figure 6. Since more bcc crystal structure with higher microhardness formed within the alloy system, it enhances the corresponding microhardness of the HEA [40]. The addition of Al promotes the transition from fcc to bcc crystal structure in AlxFeCoNiCuCr system reported [28].

The addition of Al promotes the transition from fcc to bcc crystal structure in AlxFeCoNiCuCr system reported [28]. As the bcc is considered to be higher in microhardness than fcc structure, it shows a sharp increase when the X increases from 1.0 to 1.5. However, the addition of Al also increases the cracks. It indicates that the optimal X lies between 1.5 and 1.8 for a better balance of the microhardness and cracks.Process parameters have significant effect on the hardness. It was found the average hardness of laser-produced AlCoCrCrNi samples were approximately 13% higher than the as-cast state [29]. This is due to the grain refinement of the LENS-produced samples. It also found that the microhardness increases with the cooling rate during the laser process.

This is due to the grain refinement of the LENS-produced samples. It also found that the microhardness increases with the cooling rate during the laser process. Zhang discussed that the rapid solidification during the laser process could suppress the grain growth and hence increase the amount of grain boundaries and form the grain boundary strengthening [39].7181200Ni0Ni0.5Ni1.0Ni1.5Ni2.01100)1000V H( s s e n d r a h o r c Mi90080070060050000.20.40.60.811.21.4Distance to surface (mm)Figure 6 Microhardness of Al2CrFeCoCuTiNix high-entropy alloys using the data from [40]Compressive tests are also applied to evaluate the mechanical properties of laser-produced HEAs in Table 2. It reported the overall compressive properties of AlCoCrFeNi specimens by SEBM were superior to those casting specimens [17].

It reported the overall compressive properties of AlCoCrFeNi specimens by SEBM were superior to those casting specimens [17]. The enhanced plasticity of SEBM specimens could be ascribed to the finer microstructure and existence of fcc phases, which had higher atomic packing factor and closer packed slip planes. Post heat treatment was conducted on the DML fabricated AlCoCrFeNi samples by Wang [45].

Post heat treatment was conducted on the DML fabricated AlCoCrFeNi samples by Wang [45]. The softer fcc phase resulted in reduced yield strength with enhanced ductility and this was proved in their fractographic analysis.Table 2 Compressive properties of HEAs at room temperatureAlloy AlCoCrFeNiAlCoCrFeNi𝝈y(Gpa) 1.31 1.17 1.07 1.13 1.308±0.077 1.015±0.053 0.944±0.055𝝈max(GPa) 2.6 2.94 2.83 3.02 1.425±0.1275 1.668±0.0715 1.447±0.1358𝜀p (%) 16.8 21.6 24.9 24.2 5.6±1.9 26.4±6.7 14.5±5.3Process DML, 600°C/168h HT DML, 800°C/168h HT DML, 1000°C/168h HT DML, 1200°C/168h HT Casting 1SEBM (0°) 2SEBM (90°)1 parallel to laser build direction, 2 perpendicular to laser build direction.7191.6Ref. [45] [45] [45] [45] [17] [17] [17]Corrosion ResistanceThe corrosion features of HEAs is determined by the interaction between the environment and materials [20], [46]–[48]. The common service environments include the salt water or acid.

The common service environments include the salt water or acid. The corrosion-resistance of HEAs are usually affected by the alloying element, phase change, phase segregation. Some recent working concerning on corrosion-resistance of HEAs will be discussed here.Wu et al. investigated the effect of Ni content in the alloy of FeCoCrAlCuNix and Al2CrFeCoCuTiNix coatings by laser cladding [20], [40]. SEM results showed the coating structure was consisted of heat-affected, bounding and cladding zones. Simple microstructure of fcc and bcc phases with columnar crystals were observed, which decreased the ingredient segregation. The corrosion current reduced 1 magnitude compared to Q235 carbon steel. Al2CrFeCoCuTiNi1.0 exhibited the best resistance to corrosion in the tests. However, with the increase of Ni content, the corrosion resistance enhanced first and then dropped. Wu explained that the Ni has relatively small atomic size and its lattice was seriously distorted at high contents [20].

Wu explained that the Ni has relatively small atomic size and its lattice was seriously distorted at high contents [20]. Ye et al. studied the corrosion properties of AlxFeCoNiCuCr coatings on 314L stainless steel [28], [38]. The coating had good metallurgical bonding with 314L stainless steel substrate. The relative uniform element distribution within the coatings showed better corrosion resistance than 314 L stainless steel in 0.05M HCl solution, which could be attributed to the additional of Al element.The process parameters like laser scan speed also influenced the corrosion resistance of HEAs. This can be explained that the higher the scan speed, the less laser energy was absorbed by the HEAs [13]. Therefore, the fast cooling rate made the layer surface smooth and uniform and this helped the corrosion resistance of HEAs.Table 3 summaries the electrochemical parameters through the potentiodynamic polarization tests for HEAs in aqueous environments at room temperature.

The thermodynamic theory and current additive manufacturing methods to fabricate HEAs have been reviewed in this paper. Due to the advantages of additive manufacturing, HEAs possess superior mechanical and corrosion- resistance properties to the conventional alloys. The selection of alloy element and controlling the process parameters are relatively new research topics that hold promise of future HEAs research.AcknowledgementThis project was supported by National Science Foundation Grants CMMI-1547042 and CMMI 1625736, and the Intelligent Systems Center, Center for Aerospace Manufacturing Technologies, and Material Research Center at Missouri S&T. Their financial support is greatly appreciated.Reference[1] B. J. Yeh et al., “Nanostructured High-Entropy Alloys with Multiple Principal Elements :[2][3]Novel Alloy Design Concepts and Outcomes **,” no. 5, pp. 299–303, 2004. J. Yeh, Y. Chen, S. Lin, and S. Chen, “High-Entropy Alloys – A New Era of Exploitation,” vol. 560, pp.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAdditive Manufacturing of High-Gamma Prime Nickel-Based Superalloys through Selective Laser Melting (SLM)Amrita Basak Department of Mechanical Engineering Pennsylvania State University, University Park, PA 16802, USA, Email: aub1526@psu.eduAbstractHigh-(cid:74)(cid:99) nickel-based superalloys are abundantly used in the aerospace, marine, nuclear, and chemical industries where excellent corrosion and oxidation resistance, superior mechanical properties, and exceptional high-temperature performance are required. However, selective laser melting (SLM)-based additive manufacturing (AM) of high-(cid:74)(cid:99) nickel-based superalloys pose significant challenges due to these alloys’ complex chemistry.

However, selective laser melting (SLM)-based additive manufacturing (AM) of high-(cid:74)(cid:99) nickel-based superalloys pose significant challenges due to these alloys’ complex chemistry. With multiple alloying elements and high aluminum + titanium fraction, these materials when consolidated through SLM form various secondary phases severely affecting the processability leading to the formation of cracks. The objective of this review is to summarize the progress made so far on SLM of high-(cid:74)(cid:99) nickel- based superalloys with a special emphasis towards elucidating the relationships between processing, microstructures, and properties in this alloy system.Keywords: High-(cid:74)(cid:99) Nickel-Based Superalloys, Additive Manufacturing, Selective Laser Melting (SLM), Processing, Microstructure, Mechanical Properties1.

IntroductionHigh-(cid:74)(cid:99) nickel-based superalloys are abundantly used in aircraft and power generation turbines, rocket engines, and other high-temperature environments because of their excellent corrosion and oxidation resistance, and superior mechanical properties [1]. Higher thermal gradients, thermal cycling and strains in gas turbine hot-section components cause severe fatigue loading, leading to fatigue cracking and spallation. The overall complexity associated with through additive processing advanced hot-section high-(cid:74)(cid:99) nickel-based manufacturing (AM) originates from the fact that while these alloys can be cast into complex shapes through expensive and exotic investment casting techniques, once cast, these alloys are considered non-weldable or non-joinable. A key source of non-weldability is due to the presence of the secondary (cid:74)(cid:99)(cid:3)precipitate phases in high proportion (i.e. more than 60 vol.

A key source of non-weldability is due to the presence of the secondary (cid:74)(cid:99)(cid:3)precipitate phases in high proportion (i.e. more than 60 vol. %) in these alloys arising from combined additions of aluminum (Al) and titanium (Ti) [2, 3]. While these additions provide excellent benefits in improving overall high-temperature creep properties in new castings, they cause different types of cracking such as solidification, grain boundary liquation, strain-age, and hot tearing [1].superalloysThe presence of high-(cid:74)(cid:99)(cid:3)fraction (which is an obvious peril of Ti + Al addition) increases the likelihood of cracking during post-weld heat treatment known as “strain-age cracking” [4]. The alloys also become susceptible to ductility-dip cracking associated with the formation of grain boundary carbides [5]. Based upon the compositions of %Al + 0.84 %Ti (in wt. %) versus 0.28 %Cr + 0.043 %Co (in wt.

Based upon the compositions of %Al + 0.84 %Ti (in wt. %) versus 0.28 %Cr + 0.043 %Co (in wt. %), the relationship between the alloy chemistry and the weld susceptibility is illustrated in Figure 1 [6]. While the alloys belonging to the “non-weldable”554category are the most difficult one to fabricate, the “difficult-to-weld” alloys offer many issues as they are susceptible to heat affected zone (HAZ) cracking which is also known as hot cracking or solidification cracking. The alloys belonging to “fairly-weldable” category offers difficulties with post weld heat treatment (PWHT) operations resulting in crack formations [7]. While all the AM processes work on the fundamental principle of building an object through layer-by-layer fabrication, each differs with respect to the material delivery as well as the material consolidation mechanisms. The two primary classifications of the metal AM are directed energy deposition (DED) and powder bed fusion (PBF) [8, 9].

The two primary classifications of the metal AM are directed energy deposition (DED) and powder bed fusion (PBF) [8, 9]. The material delivery for DED can be either via powder-blown or wire-fed mechanism and the energy source can be either a laser or an electron beam. Hence, in DED material is deposited where it is required. On the other hand, in PBF, the powder is spread in layers and selectively consolidated. Like wire fed DED, the energy source for PBF can also be either a laser or an electron beam.Selective Laser Melting (SLM) is a PBF-based AM process that that uses high intensity laser as an energy source to melt and fuse selective regions of powder, layer by layer, according to computer aided design (CAD) data. SLM has been shown to be capable of producing fully dense near net-shape components without the need for post-processing through complete melting the powder material.

SLM has been shown to be capable of producing fully dense near net-shape components without the need for post-processing through complete melting the powder material. However, like other AM processes, SLM parts are subjected to complex thermal including cycles repeated heating and cooling and due solidification, respectively. Such alternating thermal cycle creates residual stress in the part as well as heterogeneity and anisotropy in the part’s microstructures and properties [10]. SLM of high-(cid:74)(cid:99)(cid:3)nickel- based pose superalloys additional challenges due to complex these chemistry. Although, several review articles are available on AM of metallic materials and nickel-based superalloys [8-12], a focused review on the SLM-based processing of an important class of material system such high-(cid:74)(cid:99)(cid:3)nickel-based superalloys, is still missing.

The objective of this review is to summarize the progress made to date on these alloy systems with a special emphasis towards elucidating the relationships between processing, microstructures, and properties facilitating direct manufacturing of high-(cid:74)(cid:99)(cid:3)nickel-based superalloys such as CM247LC [13-18], CMSX-4®[19-23], MAR-M247 [24-26], IN100 [27, 28], IN738LC [29-36], K418 [37], René 142 [38-40], René 80 [41], and René N5 [6, 42, 43] for high-temperature applications. The review also identifies the potential areas of future research that is required to advance the field of SLM of high-performance material systems.10 9 8 7 6 5 4 3 2 1 0Non-weldableIN100Difficult-to-weld)%IN738LC. tFairly- weldableU700René 80wMAR-M247(CMSX-4®K418 CM247LC] iIN738meltingtoRR1000T %René N5IN939WaspaloyRené 142[ 4 8.

tFairly- weldableU700René 80wMAR-M247(CMSX-4®K418 CM247LC] iIN738meltingtoRR1000T %René N5IN939WaspaloyRené 142[ 4 8 . 0 + ] lRené 41Haynes 282Readily-weldableIncreased aging responseA %Hastealloy-XIN718[IN62556074 0.28 [%Cr] + 0.043 [%Co] (wt. %)2318alloys’Figure 1. Weldability assessment superalloys (adopted from [6]).for nickel-based5552. Selective Laser Melting (SLM)SLM starts with a CAD model of the part that is sliced at discrete intervals in the build direction. The resulting cross-sections are stored sequentially in a “build file”. Scanning algorithms are devised to plan the path of the laser beam. The powder bed is sintered conforming to the geometry requirement for the layer [12]. Next, the power bed is lowered by a distance equal to the layer thickness and a fresh layer of powder is deposited by a roller or cassette mechanism.

Next, the power bed is lowered by a distance equal to the layer thickness and a fresh layer of powder is deposited by a roller or cassette mechanism. The next cross-section is then laser- fused to the bind continuously underlying layer, and thus, by fusing layer upon layer, the entire part is fabricated. The schematic of the process is shown in Figure 2. The components created through direct SLM processes are fully dense or nearly fully dense and can be post- isostatic processed pressing full density through the closure of any residual porosity.LaserScannerdeeplyenough toPowder bedit~~~~Ir-Powder ll F~bricationdel!very pistonpistonusinghot(HIP)to achieveFigure 2. Overview of the SLM process (adopted from [8]).High power CO2, Nd:YAG, and Yttrium lasers are typically used as the energy source in SLM. The average layer thickness for SLM ranges from 20 (cid:80)m – 100 (cid:80)m for multi-pass fabrication [44].

The average layer thickness for SLM ranges from 20 (cid:80)m – 100 (cid:80)m for multi-pass fabrication [44]. For single-pass repairs, the powder layer thickness can go up to 2000 (cid:80)m [19]. The laser beam can go up to scan speed of 15 m/s. The typical spot size ranges from 40 (cid:80)m – 180 (cid:80)m in diameter [45]. SLM can be conducted at extremely pure inert environment with less than 10 PPM O2 level in order to minimize oxidation. The powder bed can be kept at atmospheric or elevated temperatures depending on the application requirement. The SLM system has the flexibility to be operated with open loop or closed loop control schemes facilitating a tighter control of the melt pool temperature [12] [46]. In SLM, preheat temperature in the range of 90 – 1200 oC has been reported in literature for processing high-(cid:74)(cid:99)(cid:3)nickel-based superalloys [26].

In SLM, preheat temperature in the range of 90 – 1200 oC has been reported in literature for processing high-(cid:74)(cid:99)(cid:3)nickel-based superalloys [26]. Scanning laser epitaxy (SLE) deserves a special mention in this regard as this process does not preheat the powder bed at all [19, 20].The parts fabricated through SLM have higher dimensional accuracy and surface smoothness than DED. Since, the part is immersed in the powder bed; this process also has advantages as the bed acts as a support structure. However, since this process requires a controlled manufacturing atmosphere, the size of build chamber, and therefore, the part is limited. Like DED or EBM processes, due to rapid alternating thermal cycles, the parts may develop residual stresses requiring further post processing. Laser powder bed fusion processes have high cooling rates, varying between ~104 and 106 Ks−1.

Laser powder bed fusion processes have high cooling rates, varying between ~104 and 106 Ks−1. On the other hand, processes like EBM use a hot bed (>870 K) reducing the cooling rate and hence do not produce a fine microstructure like the laser powder bed fusion process [47]. Since, the SLM involves partial or556complete melting of previous layer(s), the grain structure in the current layer is controlled by that of the previous layer.3. Nickel-Based SuperalloysNickel-based superalloys generally contain several different alloying elements to attain the desired property level for hot rupture strength, fatigue resistance, and creep strength. Typically, high-gradient casting (investment casting with high and well controlled thermal gradient) is used to develop customized and fine structured superalloys, although some wrought superalloys are also in use.

Properties of the superalloys are controlled by the presence of different phases in the microstructures such as the close packed phases ((cid:74)(cid:99), (cid:74)(cid:99)(cid:99), and η etc.) and the carbides (MC, M6C, M23C6 where M denotes the metal) [1]. Nickel-based superalloys primarily consist of intermetallic precipitates in a solid solution of face-centered cubic (FCC)-Ni matrix or (cid:74) matrix. The intermetallic precipitates are known as (cid:74)(cid:99), and a high-volume fraction of (cid:74)(cid:99) (up to 70%) is typically added to improve the strength of the γ-matrix. The(cid:3)(cid:74)(cid:99)(cid:3)phase is chemically represented as Ni3(Al, Ti), with Al and Ti replacing the corner Ni atoms in an FCC structure and forming a coherent phase with FCC-L12 phase. The secondary phase precipitates retard dislocation motions by forming anti-phase boundaries and locking the dislocations at high temperatures [48, 49].Table 1.

The secondary phase precipitates retard dislocation motions by forming anti-phase boundaries and locking the dislocations at high temperatures [48, 49].Table 1. Nominal composition of high-(cid:74)(cid:74)(cid:99)(cid:3)nickel-based superalloys that are processedusing SLM to date (wt.

The μ phase is typically of rhombohedral crystal structure and is found in alloys with excess percentage of Mo or W. The (cid:86) phase is a tetragonal phase that forms after prolonged exposure to a temperature between 540 oC and 9805577.1 8 2 3 5 6 - 7 - - - - - - BaloC. The hexagonal Laves phase is forms due to high-temperature exposure. These phases are topologically close-packed (TCP) phases and cause reduction in rupture strength and ductility [50]. Table 1 documents the nominal composition of high-(cid:74)(cid:99)(cid:3)nickel-based superalloys that are processed using SLM to date.Different forms of cracking are present in SLM of nickel-based superalloy components. They can be primarily distinguished as solidification cracking, grain boundary liquation cracking/hot tearing, strain-age cracking (cold cracking) and ductility dip cracking [51].

They can be primarily distinguished as solidification cracking, grain boundary liquation cracking/hot tearing, strain-age cracking (cold cracking) and ductility dip cracking [51]. Solidification cracking occurs when the mushy, two-phase liquid-solid region experiences tensile stresses and the high fraction of solid present (typically solid fraction > 0.9) restricts the flow of liquid metal to backfill the interdendritic regions [52]. The interdendritic regions are torn apart by tensile thermal stresses generated behind the melt pool as it progresses. Formation of solidification cracking is promoted by a wide solidification range for the alloy and low scan speeds. Local or partial dissolution of grain boundary phases leads to the formation of low melting point eutectics and melting of grain boundary regions, thus, leading to grain boundary liquation cracking or hot tearing when the liquid film in HAZ fails under tensile stress.

Typically, the intermetallic carbide compounds (MC and M6C) and different secondary phases (Laves- and σ-phases) contribute to partial dissolution, and hence, reduction of the melting point. Different impurities such as sulfur, boron, and phosphorus act to suppress the eutectic melting temperature and consequently increase the HAZ fissuring susceptibility [1]. A second mechanism known as constitutional liquation particular to precipitation-hardened materials occurs when a constituent particle such as TiC reacts with the surrounding matrix causing localized melting. Creation of smaller grains exposes more boundary area and lowers the strain value, resulting in reduction of hot tearing. It is also verified that processing conditions such as heating rate affect the formation of this kind of cracks. Higher power results in more area being liquated increasing the tendency of HAZ cracking [1].4.

Higher power results in more area being liquated increasing the tendency of HAZ cracking [1].4. Effects of Powder PropertiesDue to their much higher surface area per unit volume compared to cast or wrought component, powder metallurgy-based AM processes such as SLM typically have a higher degree of contamination. These contaminants adversely affect the mechanical properties of the final parts [53]. Table 2 lists the powder material properties that can significantly affect the properties of the SLM fabricate components [54, 55]. Powders used in SLM are produced via plasma, gas, or water atomization processes. In gas atomization, the molten material is atomized by inert gas jets and the resulting metal droplets are cooled in the atomizing tower. Metal powders obtained by gas-atomization typically show spherical particles. The powder quality depends on the properties of the atomizing gas such as type and pressure.

Metal powders obtained by gas-atomization typically show spherical particles. The powder quality depends on the properties of the atomizing gas such as type and pressure. The molten metal is typically produced by vacuum induction melting or electron induction melting. The gas-atomization method also works for both reactive and non-reactive material systems. In plasma atomization, metal wires that are fed into the tip of multiple plasmas, melt and are atomized in a single step. In water atomizing, a falling stream of molten metal is impinged with jets of water which immediately solidify the metal into granules (>1 mm) or powder (<1 mm). The atomization processes critically affect the powder properties such as particle size, morphology, and chemical composition impacting the final part properties. Table 3 documents the powder properties of several high-(cid:74)(cid:99)(cid:3)nickel-based superalloys that are processed using SLM to date.558Table 3.

SEM images of the powder batches showing particle morphology (adoptedfrom [34]).Engeli et al. processed several batches of commercially available IN738LC powder made by gas and water atomization under identical processing conditions [34]. Tables 4 and 5 list the powder composition. All powders were gas atomized (GA) except for batch F which was water atomized (WA). Powder batches D, E, and H were atomized from vacuum molten feedstock material (VMF), whereas batch C was directly alloyed from the elemental raw materials. No information was available for powder batches A, B, F, and G. To assess the flowability, the Hausner ratio, defined as the ratio between the tapped and apparent density, was used. A low Hausner ratio represents a good flowability. Figure 3 shows the SEM images of the power particles.

A low Hausner ratio represents a good flowability. Figure 3 shows the SEM images of the power particles. A striking difference may be observed for WA batch F that does not show any spherical powder particles at all.44.34.24.1F ,__,4 3Apparent density [glccm]2Al1C .]: -~ 0.5 1--- - - - - ---i - - - - - - - - - --i~,, -l tl - -- ·-·--·--------1.21.51.11.3 1.4 Hausner ratio [·I1.6Figure 4. Porosity the SLM processed sample versus the Hausner the corresponding powder ratio of batch. The apparent density of the powders is color coded in grayscale, with exception of batch F which is out of bound of the applied color scale (adopted from [34]).inEngeli et al. reported specimens with a density > 98.2% for all GA powder batches [34]. The process parameters optimized for powder batch H, were suitable to process all GA powder batches despite the differences in their particle size distribution and composition.

The process parameters optimized for powder batch H, were suitable to process all GA powder batches despite the differences in their particle size distribution and composition. Batches A and C showed higher porosity due to their high Hausner ratio. A low flowability can result in a local recoating. The starvation of powder after correlation between the porosity and the Hausner ratio of the corresponding powder batch is shown in Figure 4. The poor flow behavior of the powder batches F, A, and C resulted in the561lowest measured apparent density of the specimens. WA powder batch showed a considerably higher due to poor flowability resulting in poor recoating behavior. Another potential reason is due to the higher oxygen content of WA powder that results in increased oxidation and decreased wettability behavior of the melt pool.Engeli et al. reported a large variation of crack densities ranging from 0.3 mm/mm2 up to 4.5 mm/mm2 different batches of IN738LC [34].

reported a large variation of crack densities ranging from 0.3 mm/mm2 up to 4.5 mm/mm2 different batches of IN738LC [34]. The best linear correlation with the crack density was found with Si, followed by the Ga and Pb content as shown in Figure 5. Risse et al. showed that the crack density in SLM processed IN738LC was influenced by the melt pool size, however, no correlation between the melt pool size and crack density was reported [56]. With respect to the composition, in conventional welding, certain elements have been shown to increase susceptibility to cracking [57]. Rappaz et al. introduced the concept of “attractive” and “repulsive” grain boundaries [58]. The authors proposed that two approaching grains would coalesce when the energy required to create a grain boundary is lower than the energy of two solid–liquid interfaces.

The authors proposed that two approaching grains would coalesce when the energy required to create a grain boundary is lower than the energy of two solid–liquid interfaces. On the contrary, a repulsive grain boundary would have a thin liquid film between the grains until a certain undercooling is reached, which increases the susceptibility to hot cracking. Potential changes of the grain boundary or solid-liquid energies by Si would promote cracking during solidification. Zhu et al. reported that the segregation of Si increased with increasing solidification velocity. The effect of Si would therefore be amplified by the rapid solidification during SLM [59].5NE 4i.§. 3 ~"' C .g 2ii It!!!tt" u J 0 0!! U 10.1 Si [wt-% ]0 .150 .2t •5! t10Ga [wt-ppm)150515 Pb [wt-ppm)10Figure 5. Crack density measured for the different powder batches plotted versus their Si,Ga, and Pb content (adopted from [34]).5.

Crack density measured for the different powder batches plotted versus their Si,Ga, and Pb content (adopted from [34]).5. Effects of Processing ConditionsCharacteristics of SLM fabricated deposits are a strong function of the processing conditions as well as process history. The cooling rates are extremely high (~106 K/s) in SLM. Basak et al. reported the single pass fabrication of CMSX-4® specimens using SLM [19]. The authors used an additional parameter, number of repeat scans (N) to establish a steady melt pool in the beginning of the scan. N was defined as the number of raster movements that the laser performed at the leading edge of the specimen without moving in the forward direction during the start of the scan process. Reduced N resulted in discontinuous substrate-powder metallurgical bond; however, increased N caused excessive melt back near the starting edge.

Reduced N resulted in discontinuous substrate-powder metallurgical bond; however, increased N caused excessive melt back near the starting edge. Similarly, if the56220scan speed was too high, there would be a thermal lag, resulting in deformation or warping of the sample as shown in Figure 6.(a)(c)(b)(d)Figure 6. Representative transverse optical micrograph of CMSX-4® deposit showing (a) crack-free and dense deposit, (b) lack of fusion due to reduced number of repeat scans, (c) deformation/warping due to higher scan speed, and (d) lower single crystal height due to lower scan spacing (adopted from [19]).A. DensificationPerevoshchikova et al. used Doehlert matrices to investigate the influence of the three speed, processing parameters power, and hatching distance) to arrive at a set of optimal parameters to meet the component densification objective [33].

The experimental field was defined as power (P) ranging from 270-370 W, scan speed (V) ranging from 850-2250 mm/s, and hatching spacing (h) 0.09- 0.13 mm. The result of Doehlert design with three parameters was a second- order equation with one constant term, interaction three terms and three-square terms: Y = a0 + 2 + a1X1 + a2X2 + a3X3 + a11X1 2 + a12X1X2 + a13X1X3 + a23X2X3. a33X3 Here X1, X2, and X3 were coded factors, and Y was response. Coefficients a0 . . . a23 were evaluated from Y and X1, X2, and X3 using: a = (Xt X)-1XtY where Xt = (X)T. In this study, relative density of sample, (cid:85) the (defined as the ratio of specimen density over theoretical density) was the evaluation indicator of response Y, considering the influence of three variables, power (X1), scan speed (X2) and hatching distance (X3). The relative density as a function of energy density of SLM-processed(scan700al :.

The relative density significantly reduced with an increase in V and h, and decreased with an increase in P.B. Deposit CharacteristicsAcharya et al. implemented Phase 1 ascent design of experiments (DoE) methodology based on a fractional factorial design to evaluate the effects of P, V, and N on the deposit characteristics such as deposit uniformity, melt depth, and deposit height for René 80 [41]. The explored range of parameters included 100 – 400 repeat scans, 450 – 650 mm/s scan speed, 12.7 mm hatch spacing, and 350 – 600 W laser power. The deposit uniformity was calculated by taking the deposit height of the specimen, developing a linear fit for the height and then calculating the average R2 residual between the height and the linear fit. By definition, higher R2 value implied more uniform deposit. P and V were both used to control the amount of energy being applied to the specimen during the fabrication process.

By definition, higher R2 value implied more uniform deposit. P and V were both used to control the amount of energy being applied to the specimen during the fabrication process. N affected the formation of a melt pool at the start of the scan. Hence, an optimized value of N was required in order to establish a melt pool that propagated steadily fusing powder to the underlying layer. The authors reported that P impacted the deposit height and melt depth and increased each of these values as it increased (Figure 7). V impacted the melt depth and caused it to decrease as V increased. N impacted the deposit uniformity and melt depth, increasing both as N increased.C. MicrostructureMicrostructure of SLM deposited nickel-based superalloys typically show columnar, equiaxed or mixed microstructure depending on the thermal conditions and the cooling rates.

MicrostructureMicrostructure of SLM deposited nickel-based superalloys typically show columnar, equiaxed or mixed microstructure depending on the thermal conditions and the cooling rates. While an equiaxed microstructure can be size and characterized by distribution; a columnar microstructure requires several additional parameters such the primary dendrite arm spacing as columnar-to-equiaxed (PDAS), oriented-to- transition misoriented the presence of stray grains (SG) for accurate representation. An in PDAS adversely affects the high-cycle fatigue life [19]. During solidification, the solidified the region acts as a heat sink, and of gradient temperature the ahead solidification resulting front decreases, columnar-to-equiaxed transition or CET. A second solidification phenomenon, known as the oriented-to-misoriented transition (OMT), occurs when there is a change in the direction of the temperature gradient in the solidifying metal.

A second solidification phenomenon, known as the oriented-to-misoriented transition (OMT), occurs when there is a change in the direction of the temperature gradient in the solidifying metal. Existing literature reveals that the local inhomogeneity in the temperature gradient and the solidification velocity or the presence of broken dendritic arms and subsequent convection of those broken pieces acts as sources of new35the grainy = 0.0006x3.2405 R² = 0.65073025)m (cid:80)(cid:80) ( S A D P20the (CET), transition15the (OMT), and105increase0182024 22 E (J/mm3)2628Figure 8. Effect of E on the PDAS [25].564nucleation sites that result in the formation of stray grains (SG) resulting in higher grain boundary wetting and making the material prone to cracking.Basak et al. reported single pass depositions of CMSX-4® and the authors showed SGs decreased for higher values of energy density (E) [19].

reported single pass depositions of CMSX-4® and the authors showed SGs decreased for higher values of energy density (E) [19]. Lower V and higher P caused smoother flow of powder and lowered the Marangoni velocity at the top surface of the melt pool, thus decreasing the convection of broken dendrites. This reduced stray grain formation. However, the columnar ratio decreased significantly for higher values of E. The higher laser powers and lower scan speeds expanded the melt pool, thus reducing the temperature gradient in the vertical direction. This caused nucleation of new grains due to constitutional undercooling resulting in lower columnar ratios at higher values of E. Basak et al. also reported the effects of processing conditions on PDAS for MAR-M247 as shown in Figure 8 [25].

Basak et al. also reported the effects of processing conditions on PDAS for MAR-M247 as shown in Figure 8 [25]. However, no current literature is available on the effects of processing conditions on precipitate size, composition or elemental segregation in the SLM fabricated nickel-based superalloys.a)D.Metallurgical DefectsSeveral defects such as cracks, porosity, and oxide inclusions are found in the SLM fabricated nickel-based superalloys. Poor weldability of high-(cid:74)(cid:99) nickel-based superalloys results in the formation of cracks in the SLM manufactured specimens. A traditional method of reducing build-up of residual stresses is to divide the build area of each layer into smaller square sections referred to as ‘islands’, consisting of a border and inner raster scan region. Variation in cracking density using ‘island’ scanning in CM247LC deposits was reported by Carter et al. [15, 18]. Catchpole-Smith et al.

Variation in cracking density using ‘island’ scanning in CM247LC deposits was reported by Carter et al. [15, 18]. Catchpole-Smith et al. developed novel fractal scanning strategies based on mathematical fill curves, namely the Hilbert and Peano-Gosper curves (shown in Figure 9) for reducing residual stresses in SLM processed CM247LC specimens [17]. The highest XY plane bulk density of 98.0 ± 0.1% was achieved using the Peano-Gosper scan at 75 W laser power and 600 mm/s scan speed. This is significantly higher than the XY plane bulk density of the ‘island’ scan sample at 96.0 ± 0.6%. This was attributed to the large amount of spherical porosity seen in the ‘island’ sample. On the contrary, porosity was rarely seen in any of the fractal scan samples.HtJbertb)Figure 9. Illustration of the novel scan strategies generated using the (a) 3rd order Hilbert and (b) 2nd order Peano-Gosper mathematical area filling curves.

Illustration of the novel scan strategies generated using the (a) 3rd order Hilbert and (b) 2nd order Peano-Gosper mathematical area filling curves. The 100 (cid:80)m characteristic scan vector length is shown (adopted from [17]).Cloots et al. investigated the effects of Gaussian and a doughnut beam profile on the microstructure and crack formation of IN738LC [31]. Within the Gaussian series, V of 900 mm/s represented the inflection point between high and low porosity. In general, porosity within the Doughnut series was higher due to bonding defects. Porosity caused by gas pores and spherical pores was565relatively constant at 0.2% over the entire range of V. The crack density [mm/mm²] defined as the ratio of the cumulated length of all cracks detected on a certain surface of each investigated sample was about 0.27 mm/mm² for Gaussian series.

The crack density [mm/mm²] defined as the ratio of the cumulated length of all cracks detected on a certain surface of each investigated sample was about 0.27 mm/mm² for Gaussian series. In contrast, the crack density within the Doughnut series at the same scan speed was only 0.11 mm/mm², thus, establishing benefits of the Doughnut series.Qiu et al. showed the dependence of defects (pores and cracks) on laser processing parameters such as P and V [32]. With increased V, the pore size increased, and their shape became elongated due to lack of fusion. The crack density, in contrast, remained relatively constant with P within 200 – 360 W range. This was in sharp contrast with results on CM247LC that showed significantly high density of cracks [17, 18]. A significant number of cracks were found to extend along the grain boundaries (GBs). However, a few cracks initiated from the pores as well.

A significant number of cracks were found to extend along the grain boundaries (GBs). However, a few cracks initiated from the pores as well. Majority of cracks were found to be associated with the presence of particles or micro-segregation along GBs. Some particles were rich in Al and O implying the formation of Al2O3. Other particles were rich in W, Si, and O, indicating that they could be a mixture of W- and Si-oxides such as WO2 and SiO2. With regards to segregation, in the vicinity of the cracks; W, Si, and O were found to be segregated.(a)1200(b)120010001000'ii a. ~1/) 1/)iCl)800600400Transvesal direction Vertical direction~(U 800 a. ::!!:(JJ 600 (JJ ~ ii5400200200005101520Strain(%)2530350Figure 10. Mechanical properties of as-fabricated SLM specimens at room temperature, (a) Tensile test curves and (b) Corresponding mechanical data. (adopted from [37]).6. Mechanical Properties A. Tensile Strength Chen et al.

(adopted from [37]).6. Mechanical Properties A. Tensile Strength Chen et al. investigated the tensile properties of the SLM-processed K418 specimens in the build and scan direction as shown in Figure 10 [37]. The tensile strength and ductility in both directions were larger than the casting probably due to the formation of the ultrafine columnar grains due to the high thermal gradient and the rapid cooling rate during SLM. The longitudinal specimens (ε = 30.42 ± 1.68%) showed significantly higher elongation than the transverse ones (ε = 1.27 ± 2.14%). The tensile loads were parallel to the columnar grain-boundary in vertical specimens as the plastic deformation was easier when the columnar grains were elongated. Rickenbacher et al.

The tensile loads were parallel to the columnar grain-boundary in vertical specimens as the plastic deformation was easier when the columnar grains were elongated. Rickenbacher et al. reported that the yield and ultimate tensile strength of IN738LC specimens56640363228~24 ~ C 20 .Q iii 16 g' 0 12 W840extracted from the scan direction were both 23% higher whereas the average elongation was identical with the mean values of cast specimens [29]. Specimens extracted along the build direction also showed superior yield and ultimate tensile strength with 3 and 23% higher values, respectively. However, they showed a 21% lower Young’s modulus and a 49% higher elongation at break, indicating a more ductile material behavior in this direction. The difference in the properties was primarily attributed to the grain orientation and higher amount of transverse grain boundaries.B.

The difference in the properties was primarily attributed to the grain orientation and higher amount of transverse grain boundaries.B. Relaxation Behavior700II,, 600 i I5001..._ Cast refemnce -- sL , # 4: .z-speo.lmen ...... SLM, # 5: z-speoimen - SLM, # 2: xy-specfmen _. SLM, # 1: xy-specimen,(I) ,(I)I ~ 300 00200100~~ ~ - ~ ...... ~.--0010.2030405060 Time [h]7080,go,100110ote: A cast specimen was tested a referenceFigure 11. Relaxation behavior of SLM specimens, two built up along the z-axis and two within xy-plane (adopted from [29]).Rickenbacher et al. reported the results from the relaxation tests at 850 oC as shown in Figure 11 [29]. As the specimens have different Young’s moduli, the initial strain of 1% resulted in different stress levels. The relaxation behavior of the cast IN738LC specimens was superior compared to those manufactured by SLM.

The relaxation behavior of the cast IN738LC specimens was superior compared to those manufactured by SLM. This behavior was either due to differences in grain size/grain orientation and/or due to differences in the (cid:74)(cid:99)(cid:3)size and morphology. Another important difference between SLM and cast IN738LC specimens was the presence of the carbide precipitates. The specimens extracted from the scan direction showed inferior relaxation behavior and superior strength properties, while for the specimens extracted from the build direction showed the opposite behavior. Such behavior was probably due to fact that straining was primarily parallel to the columnar grains for the scan direction specimens while for the built direction specimens; strain/stress was applied transverse to the grain boundaries.567,__,__120C. Creep Results from 1,000 h and 3,000 h creep tests performed by Rickenbacher et al. confirmed the behavior of the relaxation tests [29].

Creep Results from 1,000 h and 3,000 h creep tests performed by Rickenbacher et al. confirmed the behavior of the relaxation tests [29]. Specimens extracted along the build direction showed superior creep properties compared to the specimens extracted from the scan direction, but slightly inferior compared to cast IN738LC as shown in Figure 12. SLM specimens showed lower creep rupture strengths. However, specimens extracted from the build direction were in the lower scatter band of the cast specimen due the presence of columnar grains along the testing direction. Additionally, the authors reported that the specimens extracted from the build direction approached the mean value exhibited by the cast specimens with longer test duration, whereas the specimens extracted from the scan direction tend to go astray.7.

Comparison of stress to 1% total plastic strain data of cast and SLM processed IN738LC at 850 OC (adopted from [29]).for CM247LC was effective in closing micro cracks. Messé et al. investigated the effects of stress relieving heat treatment (standard heat-treatment of 1120 °C for 2 h, followed by 845 °C for 16 h; both steps were followed by air-cooling) on as deposited (SLM), the stress relieved (SLM-SR), and conventionally cast (CC) IN738LC [30]. The (cid:74)(cid:99)(cid:3)dissolution began at 1155 °C, 1160 °C, and 1150 °C, and ended at 1175 °C, 1178 °C, and 1185 °C, for the SLM, SLM-SR and CC conditions, respectively. Though all specimens shared the same solidus temperature of 1260 °C and the same liquidus temperature of 1320 °C, the differences in the (cid:74)(cid:99)(cid:3)solvus and the range568of temperatures over which the (cid:74)(cid:99)(cid:3)dissolution happened were due to (cid:74)(cid:99)(cid:3)population difference as well as compositions. Messé et al.

Messé et al. reported that the most noticeable in the SLM-SR occurred between 1040 °C and 1090 °C. A similar event was also visible in the SLM condition, through much weaker in magnitude. The absence of such endothermic flow in CC state indicated that the transition was associated with the thermo-mechanical history.Table 6. Average sizes of the cells, grains, and precipitates of CM247LC in the as-SLM and SLM+HT (T) states. Elongated cells and grains were measured along their transverse or/and longitudinal sections, TS and LS respectively. All precipitates were measured along the TS. (adopted from [14]).ConditionCells TS ( nm)Cell colonies - grainsCarbides fraction (%vol)Carbides size(nm)'I size ( nm)15 (µm)IS (µm )as-SLM SLM + HT(1150 °C)SLM + HT c121o •q SLM + HT(1230 °()500 700700110012 13121253 5350480.80.80.850 240230250st, 50~ 15ot. 570~ 15ot. 45~ 86~.

570~ 15ot. 45~ 86~. 290", 225• 210 •SLM +liT/A701751.1270(1240 °()SLM + HTN/A802001.5210150°( l260 °C)No recrystallised grain.s: t irregular small. :t irreg ul ar large, c c uboidal. Re- crystallised grains: ♦ cuboidal.that treatment increasing temperature to 1260 °C for CM247LC resulted in a reduction of the (cid:74)(cid:99) precipitate size to approximately 150 nm as shown in Table 6 [14]. Hardness tests revealed that as-SLM specimens showed hardness of 409 ± 7 HV which continuously increased to 442.7 ± 16 HV for HT of 1210 °C, 437.7 ± 19 HV for HT of 1230 °C, 448.7 ± 23 HV for HT 1240 °C, and 462.7 ± 13 HV for HT 1260 °C. The authors acquired spectroscopy the (RUS) spectra from the specimens to extract the elastic stiffness coefficients, shear moduli and Young's moduli, Poisson's isotropic, assuming ratio, hexagonal, and orthorhombic average symmetries.

Figure 13 shows the elastic anisotropy factors obtained from the as-SLM and SLM+HT samples. For the as-SLM sample, and those solution treated at temperatures equal to or below 1230 °C, the three anisotropy factors obtained differed significantly from unity but were very similar to one another. These results confirmed that strong axisymmetric elastic bulk behavior existed in the as-SLM. Notably, the localized recrystallization observed in the SLM+HT (1230 °C) specimen did not result in a significant reduction in the elastic anisotropy. In contrast, the anisotropy factors of the specimens heat-treated at temperatures equal to or greater than 1240 °C were closer to unity. This indicated that the specimens subjected to the higher temperature heat treatments exhibited more isotropic behavior.

This indicated that the specimens subjected to the higher temperature heat treatments exhibited more isotropic behavior. Nevertheless, the deviations away569from unity in the anisotropy factors and the large RMS errors obtained when fitting the data assuming isotropic symmetry indicated that the samples heat-treated at 1240 °C and above did retain appreciable elastic anisotropy.Rickenbacher et al. reported that at elevated temperature, specimens extracted from the scan direction showed 15% higher yield strength and 9% higher Young’s modulus. The ultimate tensile strength was like cast specimens, whereas the elongation at break is 20% lower. The■■1.2!I.,,,0 u m ~ >,. C. 1.0 0.IL.. - ti) --C -dt 0.-BC.......■ A .■.. •0.16■ A • Al ... A3as-SLM 8LIM-HT SLM-HIT SLMI-HT SlM--HT 2301 iic 1240 1C 12160 tc121 o "CFigure 13.

C. 1.0 0.IL.. - ti) --C -dt 0.-BC.......■ A .■.. •0.16■ A • Al ... A3as-SLM 8LIM-HT SLM-HIT SLMI-HT SlM--HT 2301 iic 1240 1C 12160 tc121 o "CFigure 13. Elastic anisotropy factors of the as-SLM and SLM+HT CM247LC samples (adopted from [14]).specimens extracted from the build direction showed a 24% lower Young’s modulus and a 42% higher elongation in average. The yield and ultimate strength values were 5 and 3% lower compared to cast specimens, respectively.8. ConclusionsA review of the existing literature revealed that SLM of high-(cid:74)(cid:99) nickel-based superalloys has not been extensively investigated to explore and exploit the possibilities SLM offers for570fabricating high quality, low cost, repeatable, and reliable parts in automotive, aerospace, and aircraft industries.

To date, the effects of the SLM process parameters and powder properties on the physical, mechanical, and microstructural behavior of high-(cid:74)(cid:99) nickel-based superalloys under dynamic and coupled loading, oxidation, corrosion, and thermal conditions are yet to be fully proven. However, high-(cid:74)(cid:99) nickel-based superalloys are known to be uniquely challenging because of their propensity to several types of crack formation, and therefore, tend to exhibit processing problems and metallurgical defects. The current lack of thorough understanding has been reviewed to provide a basis for further research on this important class of material system.Consolidation regimes associated with SLM of high-(cid:74)(cid:99) nickel-based superalloys have been identified for a handful of alloy system. Available literature was evaluated with a view to gaining useful insights especially in the aspect of alloying effects on the microstructure and the mechanical properties.

Thus, these lightweight elements contain both geometrically simple (cover sheets) and geometrically complex (core structure) elements. Conventional manufacturing of core structures is limited in terms of geometrical freedom. On the other hand, Additive Manufacturing of sheets has disadvantages in terms of economic efficiency. Therefore, a combined process consisting of additive and conventional cost-efficient manufacturing is proposed to eliminate both disadvantages. This publication presents a hybrid manufacturing route to produce metal sandwich sheets. The hybrid sandwich sheets are manufactured using a rolled cover sheet as a base plate and additive manufactured core structures including an upper cover sheet.

The hybrid sandwich sheets are manufactured using a rolled cover sheet as a base plate and additive manufactured core structures including an upper cover sheet. For this purpose, a recently developed sheet mounting system for implementation in a laser powder bed fusion process is presented and evaluated concerning manufacturing criteria such as process stability and dimensional accuracy of the final components.IntroductionAdditive manufacturing (AM) is gaining acceptance in industry and research. Several AM processes are entering various industries and can provide a serious alternative to existing manufacturing processes. However, conventional production methods are in most cases more efficient in terms of mass production or cost-effective production chains. A combination of both, conventional and additive methods may therefore enable the avoidance of respective restrictions and the creation of synergistic effects. A promising application for this combination are sandwich sheets.

A promising application for this combination are sandwich sheets. Sandwich sheets offer a high potential for lightweight applications, such as in aircraft construction, aerospace, and the automotive industry. Further benefits may arise due to the optimal use of the material, the geometric freedom, and the possibility of integrating additional functionalities. [1] In both, purely additive and conventional manufacturing, the production of sandwich sheets currently faces challenges such as poor material utilization or long production times. In the field of AM hybrid parts, most efforts focus on fabrication based on pre-milled blanks or the realization of functionally graded materials. Yin et al. [2] demonstrated a hybrid manufacturing technology combining laser powder bed fusion of metals (PBF-LB/M) and cold spraying to fabricate multi-material parts of Al/Al+Al2O3 and Ti6Al4V.

Yin et al. [2] demonstrated a hybrid manufacturing technology combining laser powder bed fusion of metals (PBF-LB/M) and cold spraying to fabricate multi-material parts of Al/Al+Al2O3 and Ti6Al4V. Here, Al and Al+Al2O3 were cold sprayed on PBF-LB/M manufactured Ti6Al4V parts resulting in dense and machinable functionally graded materials. Although the PBF-LB/M Ti6Al4V section was represented by large pores and the cold sprayed section by small pores caused by insufficient particle plastic deformation. Tan et al. [3] investigated the interfacial characteristics and mechanical performance of steel-copper parts fabricated by PBF-LB/M. By applying appropriate laser parameters, a metallurgical bonded and dense steel-copper bi-metal was created.

By applying appropriate laser parameters, a metallurgical bonded and dense steel-copper bi-metal was created. According to the authors, the high thermal conductivity of copper led to high Marangoni convection in the interface of the materials, which in turn led to excellent bonding.1706The process concept of AM on metal sheets was already discussed in [4–8]. Ahuja et al. [5] presented a clamping concept that enabled PBF-LB/M on pre-formed metal sheets inside an SLM 280 HL machine. The mounting of the formed sheet was realized through a combination of a punch and clamping plate, which allowed the build-up on sheets with a thickness of 1.5 mm. The authors described the significant influence of an uneven sheet surface on the resulting shear strength of the hybrid components. Due to the uneven surface, an inhomogeneous powder distribution can occur on the metal sheet, which leads to insufficient bonding at the interface.

Due to the uneven surface, an inhomogeneous powder distribution can occur on the metal sheet, which leads to insufficient bonding at the interface. Although implying great potential regarding lightweight applications, the build-up of sandwich structures on thin sheets of less than 1 mm has not yet been investigated. In Rosenthal et. al [9], a new process combination route for AM sandwich sheets has been demonstrated. This work describes the AM process of sandwich sheets and their subsequent forming. As a further approach, this contribution will examine the additive build-up of sandwich sheets based on rolled sheets of high formability to develop an alternative process route that combines the high formability of rolled steel sheets with the geometric freedom of AM. To consider the lightweight aspect of sandwich sheets, we demonstrate the additive build-up on sheet thicknesses of 0.5 mm.Materials and MethodsWithin this work, 17-4 PH stainless steel powder by eos GmbH was used.

Furthermore, rolled DC01 (Fe99.19Mn0.6 C0.12 P0.045 S0.045) sheets with a thickness of 0.5 mm were applied as a base for the additive build-up. PBF-LB/M processing of the hybrid sandwich sheets was performed using an M270 system by eos GmbH, equipped with a 200 W Yb-fibre Laser (λ = 1064 nm). The utilized parameter sets, all as chessboard strategy, are shown in Table 1. The reference parameter (Ref) is the exposure parameter recommended by the machine manufacturer with a square size of 2 mm, as short scan vector lengths can lead to reduced thermal stress [10,11]. The purpose of the Ev- parameter was to reduce the temperature level that reaches the rolled sheet. Therefore, the square size was increased from 2 mm to 5 mm according to [12] to reduce the resulting temperature. Furthermore, it has already been shown in [13–15] that a reduction of the energy input can also reduce the residual stresses.

Furthermore, it has already been shown in [13–15] that a reduction of the energy input can also reduce the residual stresses. Especially a reduction of the laser power causes a beneficial decrease in the occurring residual forces [16]. Therefore, the parameter set Ev- with a reduced volume energy density of approx. 54 % was selected. Whereby the main part of the reduced energy is achieved by doubling the layer thickness and reducing the laser power by 70 W (-36 %). The line energies of the parameters are comparable resulting in a relative density above 99 % which is comparable to [17–19].Table 1: Exposure parameters used in the investigationsRef 195 laser power [W] 1000 scan speed [mm/s] 0.02 layer thickness [mm] volume energy density [J/mm³] 108.33square size [mm]2Ev- 125 700 0.04 49.60 5All experiments were executed under a nitrogen atmosphere and for all parameter sets a hatch distance of 0.09 mm was used.

Furthermore, the scan vectors were rotated in 67° steps for each layer. The build-up geometries are characterised in Table 2. The geometries are different structures of classical lightweight construction elements, but all with a thickness of the upper cover sheet of 0.3 mm. One selected structure is the honeycomb geometry (HC). The cross-sectional area1707is divided into uniform hexagonal honeycombs with a constant wall thickness of 0.18 mm. At the upper and lower end of the core geometry, holes are provided to remove the loose powder. A framework (FW) was chosen as a further core structure. The design of the geometry is x- shaped, with each element being three-dimensional. In the centre is a cross-sectional area, from which four struts extend upwards and downwards, each with a lateral offset. The single struts have a nominal thickness of 0.25 mm. The connection points are made up of four bars and therefore have a thickness of 0.5 mm.

The single struts have a nominal thickness of 0.25 mm. The connection points are made up of four bars and therefore have a thickness of 0.5 mm. Compared to the other geometries, the overall height is considerably lower at 2.48 mm. Another core geometry used is the HDP geometry. This abbreviation can be derived from the theory of the hexagonal crystal system and stands for a hexagonal densely packed lattice. The three-dimensional spheres of the same size are each connected to the surrounding spheres at four contact points. At the contact points, the contact surface is enlarged by plane sections of the respective sphere. A circular recess is created in each case, which connects the cavities and enables the removal of loose powder.Table 2: Characteristics of the investigated core structures HDP, framework, and honeycomb.

nameHDPFWHCwall thickness height sheet/core ratio0.20 mm 3.46 mm 3.520.25 mm 2.48 mm 5.830.18 mm 3.64 mm 7.18Before the manufacturing of the hybrid sandwich sheets, a mounting system is developed to ensure the successful realization of the manufacturing process. To achieve high part quality as well as to ensure the necessary process stability, the inserted sheet must be fixed plane and stable. As already mentioned above, sheet thicknesses of less than 1 mm are reasonable in the context of lightweight design. To realize these requirements (high process stability, low sheet thickness), a mounting system capable of generating a sufficient negative force in the z- direction without a form-locking connection is necessary.

Consequently, a vacuum mounting system was selected as the most suitable concept, and hereafter described in detail.Figure 1: Schematic illustration of the vacuum mounting system integrated into an eos M270 PBF-LB/M system1708A schematic illustration of the main components of the mounting system is shown in Figure 1. The system is designed for integration into the EOS M270 PBF-LB/M machine. The vacuum mounting base is located within the building chamber and is connected to the platform's lifting table through an adapter panel. Metal sheets with dimensions up to 200 mm x 200 mm can be mounted. A nozzle on the bottom of the vacuum mounting base provides a closed connection to the vacuum hose. The hose itself is led out of the system via a service tunnel, which is intended for the wiring of the conventional heating unit. At the other end of the vacuum hose, the vacuum pump is connected via a reducing adapter. The system has a total of three different filter units.

At the other end of the vacuum hose, the vacuum pump is connected via a reducing adapter. The system has a total of three different filter units. To protect the pump from powder particles that may be sucked in from the building chamber, there is a filter unit at both ends of the hose system, after the vacuum plate and before the pump. The vacuum pump used is the model E2M1.5 from Edwards Ltd. UK which is designed for continuous operation. The motor power is 0.16 kW at a constant speed of 2800 rpm with a suction volume of 1.8 m³/h and an ultimate pressure of 0.003 mbar. The required mounting area can be defined by the appropriate sealing cord (diameter 2.5 mm).Figure 2: Set-up for in-situ temperature measurements. A detailed description of the setup can be found in [20].To clarify the influence of the different process parameters on the thermally induced deformation of the rolled sheet, an in-situ temperature measurement was performed.

The setup used allowed an estimation of the thermal history in the rolled sheet and can thus provide information on the associated thermal stresses. The basic setup of the measuring system is shown in Figure 2. A detailed description of the setup is presented in [20]. Here, a C45 steel substrate plate of 220 mm x 220 mm was fixated on an adapter ring with minimal contact area. The base plate used has a recess, so that wiring for the thermocouple can be led from below the substrate plate through the shaft of the building platform. As shown, the thermocouple was spot welded directly onto the disc, ensuring maximum resolution through minimum tare volume, and thus minimal thermal inertia. A Testo 635 device with an accuracy of 0.3 °C was used to read out the data.Results The main challenge facing the manufacturing of hybrid sandwich sheets is to ensure astable additive build-up with sufficient part densities.

First, the presented part geometries (HDP, FW, and HC) were built with the Ref parameter set. All investigated core structures could be successfully built at an edge length of 35 mm x 35 mm (Figure 3a-c). When increasing one edge length of the core structure to 105 mm, the vacuum collapsed for all geometries. Resulting in a collision of the recoater blade and the part due to excessive thermal-induced deformation of the sheet. In contrast to the Ref parameter set, the utilization of the Ev- parameter set led to a complete build-up of the HC geometry (Figure 3d- f). Since only the HC structure could be built up, the core-to-sheet ratio could have a decisive role as an increased sheet/core ratio is favourable for reduced temperature input, but it might promote heat accumulation within the upper cover sheet due to insufficient heat dissipation. As the investigations of Khobzi et al.

As the investigations of Khobzi et al. [21] also demonstrated in a comparable setting.1709Nevertheless, even after a successful build-up, superelevations could be detected on one edge (Figure 3e). All other geometries (HDP, FW) considered could not be build-up either. However, upon removal of the sheet from the build chamber, severe distortion of the sheet along the y- axis was evident (Figure 3f). This distortion was compensated during build-up by the vacuum force applied in the opposite direction.Figure 3: Built-up hybrid sandwich sheet with an edge length of 35 mm x 35 mm (a-c) and 105 mm x 35 mm (d-f). Note the geometry identification in the upper left corner. d) after the build-up inside the machine, the red arrow indicates superelevations. e) full build-up hybrid sandwich sheet, the red arrow indicates superelevated edge.

d) after the build-up inside the machine, the red arrow indicates superelevations. e) full build-up hybrid sandwich sheet, the red arrow indicates superelevated edge. f) Thermal-induced deformation of the rolled sheet after the build-up, depicted by red arrows.To verify reproducibility, a total of three processes were carried out with the Ev- parameter set. After the successful build-up of the HC structure with an area of 105 x 35 mm², the resulting geometrical deviation was measured and is depicted in Figure 4. Due to the symmetry of the hybrid sandwich sheet, the ROI for the measurement of the geometric deviation is reduced to an area of about 100 mm x 15 mm to reduce the number of measuring points. The upper graph in figure 4 describes the deviation in the x-z plane, and the lower graph the deviation in the x- y plane. In the lower graph (x-y plane), the deviation of the entire area is represented by a coloured scale.

In the lower graph (x-y plane), the deviation of the entire area is represented by a coloured scale. In the upper graph, three y-positions were selected and compared.1710Figure 4: Illustration of the geometric deviation of the hybrid sandwich sheet built. Due to the symmetry of the part, the measurement was carried out on half (100 mm x 15 mm) of the sheet. Representation of the partial surface in the x-y plane in the lower graph.Figure 5: Micrographs and density analysis of the hybrid sandwich sheet. An exemplary illustration of the wall geometries with depicted nominal size in green (a), the cover sheet with depicted nominal size in green (b) and the resulting relative densities (c). (d) Etched wall geometry with the resulting penetration depth marked in red.Both graphs indicate a positive deviation from the nominal geometry at the outer edges of the sheet. This is comparable to the results in [16] for other materials.

This is comparable to the results in [16] for other materials. The maximum global deviation is about 1 mm and decreases towards the centre of the part to a minimum of 0.05 mm. With a layer thickness of 0.04 mm, this corresponds to 25 layers. It should be noted that the geometric deviation during the build-up is likely to be much smaller due to the applied vacuum force, which counteracts the induced deformation of the sheet. However, it was shown before that the parameter sets had a significant influence on the build-1711up result and even for the successful build-up with reduced energy comparably high dimensional deviations in a range of multiple layers occur. Figure 5 shows the micrographs and the corresponding relative densities of the build-up hybrid sandwich sheets. As depicted in Figure 5c, the relative density of the vertical walls is 99.27 %. An example of the wall quality is shown in Figure 5a. The edge contours are marked by process- typical sintering [22].

As depicted in Figure 5c, the relative density of the vertical walls is 99.27 %. An example of the wall quality is shown in Figure 5a. The edge contours are marked by process- typical sintering [22]. In addition, single Lack-of-Fusion defects are recognizable in the micrograph. The width of the walls is significantly wider than the nominal geometry width (0.18 mm). In addition to the surface sintering of powder particles, this may be attributable to the reduced scanning speed of the Ev- parameter, since a reduced scanning speed tends to cause a wider melt pool [23]. Nevertheless, some Lack-of-Fusion defects are still recognizable, which can be attributed to the reduced energy input [24]. The density analysis of the upper cover sheet reveals a strong geometrical deviation and, in addition, a clear expression of surface sintering in combination with a large number of defects (Figure 5b).

The significant amount of non- supported exposure area, especially during the build-up of the first cover sheet layer, led to a severe expression of surface sintering, and a predominantly uneven surface. Overall, this leads to a relative density of 97.78 % with a comparatively high standard deviation of 1.8 %. Nevertheless, the thickness of the build-up sheet (about 330 µm) is approximately in the range of the STL model of 0.3 mm. To evaluate the penetration depth of the laser and thus the quality of the metallurgical bonding between the rolled sheet and the additive structure, the etched wall geometry is shown in Figure 5d. Using the EV- parameter with 49.6 J/mm³ at 125 W laser power, the penetration depth is in the range of 50 µm. Schaub et al. [7] illustrated a penetration depth of 223 µm for a comparable volume energy density but a significantly higher laser power, 46 J/mm³ respectively 250 W.

Schaub et al. [7] illustrated a penetration depth of 223 µm for a comparable volume energy density but a significantly higher laser power, 46 J/mm³ respectively 250 W. In the same publication, they also investigated a comparable laser power of 100 W resulting in a penetration depth of 27 µm. In this context, the penetration depth of about 50 µm measured in this study is within a comparable range, which illustrates the dominant influence of the laser power.As the results of the build-up revealed, most of the process failures occurred when the upper cover sheet was reached. Only the HC core structure could be successfully built using the Ev- parameter. Consequently, the geometrical characteristics of the considered core structures seem to be crucial. Here, the core geometries differ in the varying contact areas of the core and the upper cover sheet. The difference in the sheet/core area ratio (table 2) is thereby most significant between the HDP and the HC structure.

The difference in the sheet/core area ratio (table 2) is thereby most significant between the HDP and the HC structure. In addition, [25–27] have shown that geometric features such as overhangs and the associated amount of energy directed into the loose powder influence the part temperature. A different temperature can in turn affect the deformations occurring in the rolled sheet [28]. To consider the influence of the geometry on the resulting deformations, an in-situ thermal analysis of the HC geometry (approx. 10 mm x 10 mm) was carried out using the setup introduced in [20] and [29]. The results of the in-situ thermal analysis are (exemplary for the HC structure) depicted in Figure 6. The investigated HC geometry is divided into four areas depending on the surface characteristics. The first section is characterized by an exponential increase in area.

The investigated HC geometry is divided into four areas depending on the surface characteristics. The first section is characterized by an exponential increase in area. In the second section, the exposure area is constant at 30.21 mm² until the third section shows a reduction to the level of the initial layer (14.53 mm²). The fourth sector describes the sudden increase from 14.53 mm² to 109.47 mm² in which the upper cover sheet is build-up. As the previous results have shown, this area is significant for a practicable process. Figure 6 shows the progression of the maxima for each layer relative to the initial temperature Note that the number of measurement points for the Ev- parameter set is halved due to the doubled layer thickness compared to the Ref parameter. As shown, the depicted temperature maxima increase in the first sector parallel to the exposure area per layer.

The maximum temperature difference is reached for the Ref parameter set at a z-height of 0.44 mm, respectively 0.02 mm (one layer) after reaching the z-height of the second sector (0.42 mm). For the Ev- parameter set, the highest temperature difference is present only at a z-height of 0.52 mm, thus 0.1 mm after reaching the height in the second sector. During the constant exposure area of 30.21 mm² in the second sector, the temperature differences of both parameter- sets decrease with increasing distance to the measuring position in the base plate to 5.97 K (Ref) and 3.22 K (Ev-). Due to the further successive reduction of the exposure area in the third sector to 14.46 mm², the temperature differences decline continuously. At the transition between the third and fourth sectors, there is a sudden increase in the exposure area in the sector of the upper cover sheet to 109.47 mm².

At the transition between the third and fourth sectors, there is a sudden increase in the exposure area in the sector of the upper cover sheet to 109.47 mm². In the first two layers of the upper cover sheet, downskin- parameters typical for the PBF-LB/M process and representing low energy input were used [30]. Despite the meanwhile progressed distance to the measuring position, the temperature differences resulting from the downskin-parameters can also be identified in the graph. The parameter sets Ref and Ev- used from the third layer within the fourth sector lead to a maximum temperature difference of 23.8 K (Ref) and 18.0 K (Ev-), respectively. By the time the part is finished, the temperature differences decrease again.

By the time the part is finished, the temperature differences decrease again. Based on the results, it can be deduced that the reduced temperature level during the Ev- build- up, leads to a reduction in deformation as [28] and [31] highlighted the importance of the temperature level on the resulting deformations.ConclusionAs the results of the build-up of the three different core structures and the in-situ temperature measurement of the HC structure revealed, the geometry of the sandwich sheet and1713the exposure parameters used influence the resulting temperatures and thus the forces to be compensated by the vacuum mounting system. By choosing a suitable parameter combination with reduced energy, the heat input into the rolled sheet is reduced and at the same time, heat accumulation in the upper cover sheet is avoided.

By choosing a suitable parameter combination with reduced energy, the heat input into the rolled sheet is reduced and at the same time, heat accumulation in the upper cover sheet is avoided. These results are in strict contrast to the conventional PBF-LB/M manufacturing strategies, which aim at a targeted and fast transport of the heat, especially in areas with a significant overhang [32]. The high amount of overhangs of the sandwich sheets is beneficial within the production concept presented for the manufacturing of hybrid sandwich sheets, as these overhangs protect the rolled sheet inherently from excessive thermal energy input and thus from deformation. Within further investigations, especially the influence of different core geometries should be examined by numerical studies.

Within further investigations, especially the influence of different core geometries should be examined by numerical studies. This would allow a critical sheet-to-core ratio to be determined, which in turn could be used as a design parameter in the future development of hybrid sandwich sheets.AcknowledgementThis work was kindly supported by the German Research Foundation (DFG) underproject numbers WI 2118/9-3 and WI 2118/17-1. Furthermore, we would like to thank Prof. A. Erman Tekkaya and Dr. Stephan Rosenthal from the Institute of Forming Technology and Lightweight Components at the TU Dortmund for providing the CAD files of the core structures. We also thank Fabian Schulz for the execution of the manufacturing processes within his bachelor thesis.References [1] Wohlers T, Campbell RI, Diegel O, Kowen J, Mostow N, Fidan I. Wohlers report 2022: 3D printing and additive manufacturing global state of the industry.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceAdditive Manufacturing of Liners for Shaped ChargesJason Ho1, Cody S. Lough2, Phillip Mulligan1, Edward C. Kinzel2, Catherine E. Johnson11Mining and Explosives Engineering 2Mechanical and Aerospace EngineeringMissouri University of Science and Technology, Rolla, MO 65409AbstractA Shaped Charge (SC) is an explosive device used to focus a detonation in a desired direction, and has applications in demolition and oil extraction. The focusing relies on a void in the explosive mass, shaped by a metal liner that becomes a projectile during detonation. Additive Manufacturing (AM) allows greater design freedom and geometric complexity for the liner portion of the SC. Specifically, hierarchical structuring and functional grading can potentially provide greater velocity, directionality, and efficiency.

Specifically, hierarchical structuring and functional grading can potentially provide greater velocity, directionality, and efficiency. In this work, Selective Laser Melting (SLM) is used to explore different geometries for an SC liner made out of SS 304L. These are detonated using the explosive Composition C-4 to evaluate performance metrics, depth and standoff, and are observed using high-speed imaging. The work shows the potential for advanced shaped charges produced using SLM.1. IntroductionShaped Charges (SC) have been used in various applications throughout the years including demolitions and oil well completion. The primary purpose of an SC is to penetrate barriers with a confined force. In the oil industry, the SC is used to penetrate the casing of the well and geological formations to increase oil flow.

The primary purpose of an SC is to penetrate barriers with a confined force. In the oil industry, the SC is used to penetrate the casing of the well and geological formations to increase oil flow. Charles Munroe brought hollow SC’s into the mainstream through his work in the late 1800’s where in one instance he was able to deform a steel plate in the shape of initials inscribed into a charge (explosive engraving). Munroe also found forming a conical cavity in a cylinder of explosive resulted in increased penetration into a steel target. This phenomenon became known as the Munroe effect. SC’s with metal liners were later found to have increased performance over hollow liners. [1]Birkhoff et al. demonstrated and discussed the theory behind SC’s with conical liners used for penetrating and wedge liners for cutting materials. The explosive wave generated by detonation of the charge reaches the apex of the liner. This produces high pressures and leads to a collapse of the liner.

The explosive wave generated by detonation of the charge reaches the apex of the liner. This produces high pressures and leads to a collapse of the liner. The walls of the liner begin to collapse in the direction normal to their surfaces. The inner material of the liner then forms a jet and the exterior forms a metal slug traveling at high and low velocities, respectively. [2]The shaped charge process is further discussed by Walters where it is stated that the spherical pressure waves generated by the explosion travel around 8 km/s and pressures near 200 GPa are reached during the process. Due to the extremely high pressures, the yield strength and ultimate strength of the liner material are negligible. The jet produced from the liner collapse then479generates pressures around 100-200 GPa when it strikes a steel target. This produces a cavity in the target through displacement of the impacted material. The target material displacement is caused by the extremely high pressures.

This produces a cavity in the target through displacement of the impacted material. The target material displacement is caused by the extremely high pressures. The cavity in the target is not primarily caused by a thermal process. [3]Figure 1 is demonstration of the jet formation and target penetration for a Conical Shaped Charge (CSC) through simulation. In the simulation, the condition of the materials including C4 explosive, the metal CSC liner, and the steel target are determined as a function of time after detonation. The jet and slug formed from the liner as a result of the liner collapse are highlighted in Fig 1 (b). Figure 1 (c) shows the penetration of the jet into the steel target.(a)(b)(c)C4SlugCSC LinerJetSteel TargetFIG 1: Simulation of CSC liner testing.The performance of a shaped charge is increased through optimized liner design and distance away from the target. Vigil reported on the optimized design of liners for CSC’s.

Vigil reported on the optimized design of liners for CSC’s. The optimized design for CSC liners is a function of the diameter of the cone and is included as Fig. 2. Liners are traditionally manufactured using this design by forming uniform thickness sheets of copper or steel. Copper liners have the best performance when compared to other materials.480FIG 2: Optimized design of CSC liner [4].It has been reported that liners manufactured with non-uniform thickness have the potential for increased performance. Efficient manufacturing for these liners is a challenge. Venghhiattis addressed this challenge in the 1960’s by developing a process to press powder metal into liners with non-uniform wall thicknesses [5]. Additive Manufacturing [AM] is an ideal process that can be used to efficiently produce CSC liners with non-uniform wall thicknesses, or liners with more complex designs to increase performance beyond the traditionally formed copper liners.

The goal of this paper is demonstrate the ability to additively manufacture CSC liners out of 304L stainless steel using the Selective Laser Melting (SLM) process. The hypothesis that anisotropy in the liners will affect performance is also tested.2. Experimental ApproachThe design of the CSC liners in this work was based on the optimized geometry [4] with modifications within the walls possible to manufacture with AM. The liners were manufactured using 304L stainless with a Renishaw AM250 SLM system. The SLM process parameters used during the manufacturing of the CSC liners were optimized for density. The modification to the CSC liner design included the addition of an internal honeycomb structure to alter the liner properties. Figure 3 (a) contains the CSC liner design showing the internal honeycomb structure through different slice views. An image of the completed build is included as Fig.

Figure 3 (a) contains the CSC liner design showing the internal honeycomb structure through different slice views. An image of the completed build is included as Fig. 3 (b).481(a)(b)10 mmFIG 3: (a) CSC liner design for SLM manufacturing and (b) completed build.The honeycomb structure was included in the interior of the CSC shell to alter the stiffness of the liner. It was hypothesized that changing the wall thickness of the honeycomb structure would lead to different stiffness for the CSC liners and affect the performance. Varying the wall thickness of the honeycomb also resulted in changes in liner mass. This occurred because the hollow areas of the honeycomb structure are not processed by the laser and retain the density of the trapped un- sintered 304L stainless steel powder. The laser melted areas of the honeycomb reach near full density of bulk 304L.

The laser melted areas of the honeycomb reach near full density of bulk 304L. The CSC liners were processed with keeping the center-to-center distance, d0, of the honeycomb structures a constant 2.5 mm. The width of the hollow section of the honeycomb structures, d1, was then varied. Figure 4 is the dependence of the CSC liner mass on the ratio of d1 to d0. The mass decreases with increasing d1 from the extreme case of a completely solid liner (d1 = 0 mm) to the case of a liner that is a hollow shell (d1 = 2.5 mm).10095Solid]g[s s a M9085d0Honeycombr e n Li8075d1Hollow700.00.20.4d1/d00.60.81.0FIG 4: Dependence of CSC liner mass on honeycomb parameters.The CSC liners were tested using the following procedure to observe the effects of the mass and stiffness on performance. The CSC was firm-pressed into the flat end of a PVC pipe cap, resulting in a tight seal.

The CSC was firm-pressed into the flat end of a PVC pipe cap, resulting in a tight seal. A PVC pipe was then fitted into the rounded bottom containing the CSC and sealed with rubber cement. Next 300 grams of C4 was massed and then tightly packed into the PVC pipe, covering up the CSC and filling up the tube. A rounded end of PVC pipe was rubber482cemented around the open end of the PVC tube. A detonator was inserted into a 10 gram stinger and the stinger was inserted into the hole at the top of the PVC pipe cap. A stack of 6.35 mm thick steel plates were secured to an I-beam with F-clamps to ensure they remained together during detonation. The CSC assembly was then placed at a standoff distance of 0.257 meters from the top steel plate. High speed videos were recorded during testing to determine the velocity of the jet formed for each CSC. Figure 5 (a) is a schematic of the CSC test setup Fig. 5 (b) is an image of the experimental setup.

Figure 5 (a) is a schematic of the CSC test setup Fig. 5 (b) is an image of the experimental setup. The impact site produced by a CSC in one of the 6.35 mm thick steel plates during preliminary testing can be seen in Fig. 5 (c).(a)10g Stinger(b)e s a C C V PExplosive: 300g C48 5 . 8 m m(c)CSC Liner42.9 mm2 5 7 . 4 m mS t a n d o f fSteel Plate 16 . 3 5 m m2 cmFIG 5: (a) Schematic and (b) image of the CSC test setup and (c) image of first plate impacted during trial run.3. Results and DiscussionPerformance of the CSC liners was characterized by the quantification of the jet velocity in air, the number of plates penetrated, and the number of plates deformed. Figure 6 contains selected frames from the high speed videos captured during testing with a Phantom V2012 camera. The data was collected using a frame rate of 110,000 Hz. The images in Fig.

Figure 6 contains selected frames from the high speed videos captured during testing with a Phantom V2012 camera. The data was collected using a frame rate of 110,000 Hz. The images in Fig. 6 were taken from the recordings of the tests with the hollow and solid liners at 27, 54, and 81 µs after detonation. In the recordings the metal jet generated due to the explosion is assumed to be at the bottom of the plume. With this assumption, it can be concluded that the jet for the hollow liner reached the target within 81 µs while the jet for the solid liner was still in air. This result was quantified through the calculation of the jet velocity in air for each of the five liners tested.

This result was quantified through the calculation of the jet velocity in air for each of the five liners tested. The jet velocity was determined by dividing the known standoff distance from the CSC to the first target plate by the amount of time it took the bottom of the plume to reach the surface of the plate.48327 µs54 µs81 µsHollow Liner10 cm27 µs54 µs81 µsSolid Liner10 cmFIG 6: High speed images captured during testing of hollow and solid CSC liners.The calculation results for jet velocity in air for each liner are plotted against the mass in Fig. 7. The jet velocities were determined to be on the order of kilometers per second. Figure 7 demonstrates that an increase in the CSC liner mass results in a decrease in jet velocity. The maximum jet velocity corresponds to the hollow CSC liner and the minimum jet velocity was generated using the solid liner.

The maximum jet velocity corresponds to the hollow CSC liner and the minimum jet velocity was generated using the solid liner. Assuming kinetic energy is a constant since the same amount of C4 was used in each test, a simple model for the jet velocity can be developed. In this model, the jet velocity in air is proportionate to the inverse square root of the liner mass. The jet velocities were fit to the model and the result is included in Fig. 7. The fitting results show the model deviates from the experimentally determined jet velocities. This suggests the velocity of the jet depended on more than the liner mass.

7. The fitting results show the model deviates from the experimentally determined jet velocities. This suggests the velocity of the jet depended on more than the liner mass. A possible explanation to the deviation of experimental jet velocities from the model is the anisotropy in the liner stiffness resulting from the honeycomb structures.4.25] s / m k [4.00Hollowr iA niy t i c o e Vlt e J3.753.50v = c·m-0.5HoneycombSolid3.2570758085 Liner Mass [g]9095100FIG 7: Jet velocity in air as a function of CSC liner mass.484The images in Fig. 8 are of the first two plates impacted during testing of the hollow, solid, and a representative honeycomb liner. The images for Plate 1 show each test resulted in a rough impact crater. The rough deformation of the impact zone for the first plate was most likely due to the metal slug formed at the rear of the jet and possibly some fragmentation of the CSC liner.

The rough deformation of the impact zone for the first plate was most likely due to the metal slug formed at the rear of the jet and possibly some fragmentation of the CSC liner. The first plate catches the slug and fragmented pieces of the liner. This results in a much cleaner impact craters for the subsequent plates since the leading part of jet is the only penetrating element.Hollow LinerHoneycomb LinerSolid Liner1e t a Pl2e t a Pl10 mmFIG 8: Images of impact craters for first two plates produced with hollow, honeycomb, and solid CSC liners.Figure 8 also shows each liner penetrated the first plate while only the hollow liner was able to penetrate the second plate. All three liner types were able to deform the second plate. Figure 9 contains a simple quantification for number of plates penetrated and deformed for the five CSC liners plotted against their respective jet velocities.

Figure 9 contains a simple quantification for number of plates penetrated and deformed for the five CSC liners plotted against their respective jet velocities. The lower jet velocities generated for the honeycomb structure and solid liners resulted in the penetration of the first plate. While these liners only penetrated the first plate, the deformation of the underlying plates differed. Jet velocities below 3.5 km/s deformed the first two plates. The velocities between 3.5 and 3.75 km/s deformed 3 plates. The highest jet velocity corresponding to the hollow liner penetrated 4 plates and deformed 5. An image of the 4 penetrated plates from the hollow liner test is included as Fig.

The highest jet velocity corresponding to the hollow liner penetrated 4 plates and deformed 5. An image of the 4 penetrated plates from the hollow liner test is included as Fig. 9 (c).485665(a)5(b)d e t a r t e n e P s e t a Pl432d e m r o f e D s e t a Pl432110 3.253.503.75 Jet Velocity [km/s]4.004.250 3.253.503.75 Jet Velocity [km/s]4.00(c)FIG 9: Number of plates (a) penetrated and (b) deformed as a function of jet velocity and (c) images of 4 plates penetrated by hollow CSC liner.The results in Figs. 7-9 demonstrate the hollow liner performed the best out of the CSC’s tested in this paper. This result is a combination of several factors including liner mass, anisotropy of liner stiffness, and pressure wave propagation impedance mismatch. The liners with a larger mass underperformed. The larger mass lead to smaller jet velocities and resulted fewer plates penetrated.

The liners with a larger mass underperformed. The larger mass lead to smaller jet velocities and resulted fewer plates penetrated. The anisotropy introduced into the liner stiffness also negatively affected the performance. The honeycomb structures most likely provided an increased resistance to collapsing during testing, which led to the lower jet velocities and penetration of a single plate. The primary factor explaining the differences in liner performance is the impedance mismatch during propagation of the pressure wave generated by the explosion through the CSC liner. In the CSC liners with the honeycomb structure, the impedance mismatch for the pressure wave at the interfaces of the near fully dense solidified borders and un-sintered voids resulted in energy loss. These losses were minimized for the hollow liner which only had two fully dense-powder metal interfaces contributing to the impendence mismatch.

These losses were minimized for the hollow liner which only had two fully dense-powder metal interfaces contributing to the impendence mismatch. The solid liner should have even less impedance mismatch, but the mass and stiffness were the dominating factors leading to the decreased performance when compared to the results for the hollow liner.4. Summary and ConclusionsIn this paper Conical Shaped Charge liners were additively manufactured using SLM. The liner design included an internal honeycomb structure to create anisotropy in the stiffness of the CSC liners. The wall thickness of the honeycomb structure was varied to produce five unique liners with different masses. The varied mass of the CSC liners resulted in 0.8 km/s range in jet velocities. A jet velocity of 4.2 km/s generated during testing of the hollow liner resulted in the penetration of four 6.35 mm thick steel plates. The hollow liner preformed best out of the CSC’s tested in this paper.

The hollow liner preformed best out of the CSC’s tested in this paper. The honeycomb and solid liners preformed worse due to combinations of4864.25larger masses, the negative effects of the stiffness anisotropy, and higher impedance mismatch leading to energy loss during propagation of the pressure wave.The preliminary results in this paper demonstrate the potential of using AM to manufacture CSC liners. The CSC liners were built using SLM with an off nominal material and were able to penetrate and deform steel plate targets. Future work will focus on hierarchically reducing the bending stiffness of CSC liners while increasing radial stiffness. This design will potentially lead to less resistance during the formation of the jet, which should increase the CSC performance. AM also allows the ability to tailor the material response locally by lowering yield stiffness with reduced laser energy density through process parameters adjustments.

AM also allows the ability to tailor the material response locally by lowering yield stiffness with reduced laser energy density through process parameters adjustments. The reduced restrictions for design along with the advanced manufacturing capability in AM allow the opportunity to tailor liners for peak performance using methods difficult to implement with traditional manufacturing.5. References[1] Walters W. A Brief History of Shaped Charges. 24th International Symposium on Ballistics.2008; 1:3-10.[2] Birkhoff G, MacDougall D, Pugh E, Taylor G. Explosives with Lined Cavities. Journal ofApplied Physics. 1948; 19:563-82.[3] Walters W. An Overview of the Shaped Charge Concept.[4] Vigil M. Optimized Conical Shaped Charge Design Using the SCAP Code. Sandia Report.1988.[5] Venghhiattis A. Method of Manufacturing Shaped Charge Explosive with Powdered MetalLiner. US 3255659, United States Patent Office. 1966.487.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceAdditive Manufacturing of Metal Bandpass Filters for Future Radar ReceiversBradley Grothaus1, Dane Huck1, Austin Sutton1, Ming Leu1, Ben Brown2 1Department of Mechanical and Aerospace Engineering, Missouri University of Science andTechnology, Rolla, MO 65409 2Department of Energy’s Kansas City National Security Campus Managed by HoneywellFM&T, Kansas City, MO 64147AbstractSelective laser melting (SLM) is a powder-bed fusion (PBF) process that bonds successive layers of powder with a laser to create components directly from computer-aided design (CAD) files. The additive nature of the SLM process in addition to the use of fine powders facilitates the construction of complex geometries, which has captured the attention of those involved in the design of bandpass filters for radar applications.

However, a significant drawback of SLM is its difficulty in fabricating parts with overhangs necessitating the use of support structures, which, if not removed, can greatly impact the performance of bandpass filters. Therefore, in this study bandpass filters are manufactured in two stages with 304L stainless steel where each builds only a portion of the part to improve the reliability in manufacturing the overhangs present. The results show that the versatility of SLM can produce difficult-to-manufacture bandpass filters with high dimensional accuracy. This work was funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy.1. IntroductionPowder-bed fusion (PBF) is a class of additive manufacturing (AM) techniques that refer to the selective consolidation of particles in a layer-by-layer fashion to create three-dimensional components.

Among the PBF methods is the selective laser melting (SLM) process where a laser is used to weld regions of a powder-bed to construct parts additively. With the number of increasing applications of SLM in industries including aerospace and medical fields, the amount of research being directed towards improving the process regarding manufacturability of geometrically complicated components has increased dramatically in the previous 10 years. Moreover, the types of materials manufacturable by SLM has increased to encompass primarily aluminum, titanium, nickel, and steel alloys so that the parts fabricated may be used in service.Figure 1. Illustration of overhang and dross formation [1]500Although SLM has enabled more freedom in the design process due to its ability for constructing complex geometry, this method still suffers from the difficulty in printing overhanging structures, often necessitating the use of supports that may be difficult to remove.

During the SLM process, high temperatures exist above the melting point of the metal material being processed [2]. Subsequent rapid solidification occurs which causes the formation of residual stresses, which can distort parts. Therefore, supports in SLM serve two main purposes: anchoring the solidified material to the substrate to eliminate part distortion due to residual stresses, and heat dissipation. If parts are not properly supported, the concentrated heat cannot be sufficiently dissipated through the powder bed due to its low conductivity [3], causing thermo-elastic distortion [1]. This problem is exacerbated for components that are built less than 45 degrees relative to the substrate, requiring support material to be inserted to ensure that the parts can be built [4].

This problem is exacerbated for components that are built less than 45 degrees relative to the substrate, requiring support material to be inserted to ensure that the parts can be built [4]. Under this critical angle the melt pool increases and flows into the powder bed because of gravity and both the capillary forces and thermal energy cause the underlying powder to adhere to the part surface causing warping and drastically increasing the surface roughness leading to what is more commonly known as dross formation [1], as shown in Figure 1. Many times, dross formation can cause harm to the recoater jeopardizing an entire build and the dimensional accuracy of the part.

Many times, dross formation can cause harm to the recoater jeopardizing an entire build and the dimensional accuracy of the part. Thus, these support structures become almost essential to parts built at low angles, thereby forcing geometric constraints to be applied to parts during the design process so that they may be correctly manufactured by SLM.In the case of bandpass filters for use in radar applications, these geometric constraints can greatly inhibit the applicability of SLM for their construction. However, as SLM can produce complicated geometry, the process has the potential to manufacture unique designs that are difficult to produce by conventional means. The bandpass filters of interest for this study are an interdigitated design, which are shown in Figure 2. The bandpass filters on the left and right in Figure 2 are hereby referred to as bandpass filter A and B, respectively.

The bandpass filters on the left and right in Figure 2 are hereby referred to as bandpass filter A and B, respectively. It should be noted that such both geometries are difficult to produce by conventional means as the small cantilevers can become distorted during the milling process influencing the performance of the bandpass filter, which can be avoided using SLM. Yet, the interdigitated cantilevers are still difficult to manufacture with SLM as they produce overhangs regardless of the orientation at which the parts are built. Thus, the parts pose a serious challenge in the SLM process, which serves as the basis of this paper.Figure 2. CAD models of two bandpass filter designs (not to scale).501The goal of this research is to explore and evaluate methods implementable in the SLM process for fabrication of the given parts with a Renishaw AM250 and assess each in terms of their success with respect to dimensional accuracy and overall manufacturability.

The most popular method for building these overhang structures is using supports. However, support structures can affect the SLM process in undesirable ways. The need for support structures increases the laser exposure time during the process and the removal of these structures can take a significant amount time and energy ultimately increasing the cost. Therefore, in addition to building each of the parts with supports, an alternative method is presented where only a portion of each part is built at a time to bypass the manufacture of overhanging structures, hereby referred to as “flip and fixture.” Lastly, an additional technique is used, which involved the use of non-contact supports, a common technique employed in electron beam additive manufacturing. With non-contact supports, supports are offset a certain distance from the parts to provide heat dissipation without adhering to the part surface for ease of removal.2.

With non-contact supports, supports are offset a certain distance from the parts to provide heat dissipation without adhering to the part surface for ease of removal.2. Method 1: Optimization of Supports (Unsuccessful)A common choice for difficult-to-manufacture parts in SLM is to build supporting material that will assist in fabrication. This procedure is often coupled with orienting the parts in an optimized manner that will minimize the amount of support material needed and facilitate the ease of their removal to retain the final part. Other concerns when using support material include minimization of surface roughness and dimensional accuracy, both of which are greater in SLM for overhanging surfaces due to dross formation.

Other concerns when using support material include minimization of surface roughness and dimensional accuracy, both of which are greater in SLM for overhanging surfaces due to dross formation. Thus, optimizing the support structures to mitigate all these negative effects is crucial.In the support optimization process performed in Materialise Magics, the goal was to produce structures with limited volume, small areas of connection to the part (to make the removal process easy and time-efficient) while orienting the part optimally for a good surface finish and limit down-facing surfaces. In this study there were five different parts on the support optimization build plate, each with a different set of support and orientation combinations. Elements that variedFigure 3. Method of support plate overhead view502(a)(b)(c)(d)(e)Figure 4.

Elements that variedFigure 3. Method of support plate overhead view502(a)(b)(c)(d)(e)Figure 4. Support optimization as performed in Materialise Magics for the two bandpass filters in this study: (a) line supports with orientation pertaining to minimal support surface, (b) point supports at 60 degrees from the horizontal, (c) line supports at 45 degrees from the horizontal, (d) point supports with minimal support surface and z-height, and (e) point supports with minimal support surface.with each part were the orientation and the type of support used, either a line support or a point503support. A line support is a wall of support material with perpendicular struts, while a point support simply attaches to the part at a single location. When using point supports, many were generated to simulate a “wire brush” for ease of removal.

When using point supports, many were generated to simulate a “wire brush” for ease of removal. To further enhance the ease of removal, each type of support was only welded with one laser pass making them weak due to the porosity present without remelting. To attempt different orientations for each part, the optimization tool in Materialise Magics software was utilized. This tool optimizes the orientation based off the geometry and down facing surfaces to create an orientation that will limit supports and laser exposure time. The other orientations (besides the Materialise Magics optimized orientation) used were 45 and 60 degrees. On all the parts the teeth of the supports (the very top and bottom of the supports that connect to the substrate and the component being supported) were designed to be easily removed and quick to build as was found by Calignano [5]. Also, on each cantilever there were two line supports added to support each downfacing edge of the cantilever.

Also, on each cantilever there were two line supports added to support each downfacing edge of the cantilever. The supports generated for each part can be seen in Figure 4.Part 1 was built with all line supports and the orientation was optimized to minimize the amount of support surface needed on the part (Figure 5a). This was the most successful part built. It finished building to completion and is recognizable as the part it is supposed to be. Unfortunately, the down-facing surfaces have a less than optimal surface finish and low dimensional accuracy.Part 2 was built with point supports, and the orientation was 60 degrees vertically from the horizontal build plate. This part failed early on due to the cross teeth in supports for the cantilevers overlapping. Due to this overlapping, the laser beam passed multiple times over the same section of each cantilever causing distortion.

Due to this overlapping, the laser beam passed multiple times over the same section of each cantilever causing distortion. So as not to interfere with the recoater, this meant the part had to be suppressed early in the build.Part 3 was built with line supports and the orientation was 45 degrees vertically from the horizontal build plate. This part failed as soon as the sections with a 45-degree angle without supports started building. When this occurred the part had to be suppressed.Part 4 was built with the point supports, and the orientation was optimized in the Materialise Magics surface to optimize supports and the build height. This part failed to build to completion (Figure 5b). As it was building the angle was close to the critical angle and was close to failing but the machine would correct itself before the part failed. When it reached the top, it did not correct itself therefore it had to be suppressed before completion.

When it reached the top, it did not correct itself therefore it had to be suppressed before completion. Even with most of the part that did complete, the surface finish on the down facing surfaces was not optimal.Part 5 was built with point supports, and orientation was optimized in the Materialise Magics surface to minimize the amount of support surface on the part. This part failed to build to completion. Just as Part 3 on this build plate did, it failed when it started building the opposite angled non-supported parts.3. Method 2: Non-Contact Supports (Unsuccessful)Another method that was examined in this research was the non-contact support method. The main purpose in evaluating this method an alternative means of building simple cantilever504(a)(b)Figure 5. Fabricated bandpass filters with (a) line supports and minimal support surface, and (b) point supports with minimal support surface and z-height.overhangs in the SLM process.

Fabricated bandpass filters with (a) line supports and minimal support surface, and (b) point supports with minimal support surface and z-height.overhangs in the SLM process. Unlike conventional support structures, non-contact supports are completely separated from the overhang by some gap distance. The idea is that the support will act as a heat sink and help with thermal dissipation via heat conduction. Ideally, this would lead to the overhang building without the usual distortion of the overhang occurring such as when there is no support at all. Related previous research consists of one study examining the feasibility of non-contact supports in the electron beam melting process [6]. Figure 6 illustrates the concept of the non-contact support. Note that the overhang is completely horizontal. This is the orientation of the overhangs tested throughout the method of non-contact supports.505Figure 6.

Note that the overhang is completely horizontal. This is the orientation of the overhangs tested throughout the method of non-contact supports.505Figure 6. Non-contact support with overhang CAD model.The experimental approach for testing the method of non-contact supports consists of the following: design the test setup, build the test setup, and observe the result of each test. Each test is essentially a set of cantilever overhang structures. Each test structure contains an overhang with a base and a block under the overhang. The block is separated by some gap distance from the overhang. After each test was examined, observations were drawn, and a new test was designed to make more observations.For the first test setup, the questions to consider include: the required gap distance that would be successful, if any, and whether the thickness of the non-contact support carries any effect. This led to a test setup that was used to evaluate multiple variables.

This led to a test setup that was used to evaluate multiple variables. The set of tests consisted of three different heights of overhang test structures, each having a gap distance from 0.05mm to 0.35mm (0.05mm interval jumps). These results are shown in Figure 7a. All the test structures from the previous test bonded between the overhang and the non-contact support. The second set of tests consisted of five iterations of each gap distance of 0.5mm, 1.0mm, 1.5mm, and 2.0mm. The five iterations made it possible to see any consistency or lack thereof within each gap testing and the larger gap intervals allowed for more rapid locating of a precise range for building overhangs. The test results are shown in Figure 7b.The third and final test was that of a more specific range of gap distance.

The test results are shown in Figure 7b.The third and final test was that of a more specific range of gap distance. Given that the largest three gaps of the previous test failed to build, and the smallest gap caused bonding between the overhang and the support, it appeared that the range was between 0.5mm and 1.0mm. Staying close to the smaller end of the range, and including some overlap, the next setup consisted of a similar outline. Again, five iterations of each gap distance were tested: 0.4mm, 0.5mm, 0.6mm, 0.7mm. The results are shown in Figure 7c.Three different outcomes occurred from these tests, as illustrated in Figure 8. For this method to be feasible, there needs to be another outcome between bonding and curling: building successfully without curling. Unfortunately, the non-contact supports simply do not disperse fast enough for a successful build of an overhang.

Unfortunately, the non-contact supports simply do not disperse fast enough for a successful build of an overhang. While the quality is better than the case where the horizontal overhang has no support whatsoever, it is not sufficient enough to build with dimensional accuracy and quality surface finish.5060.05mm0.10mm0.15mm0.20mm0.25mm0.30mmmmmm0.35mm(a)1.0mm0.5mm2.0mm1.5mm(b)0.5mm0.7mm0.4mm0.6mm(c)Figure 7. Non-contact support builds where (a) is the first set of tests, (b) is the second set of tests, and (c) is the third set of tests; the gap used is labeled next to each set of overhangs.Table 1 shows that 0.6 mm and 0.7 mm converge to a single curling result for each. The 0.6 mm test overhangs had four that bonded, while the 0.7 mm test overhangs had four that failed5070.6mm1.5mm0.35mmFigure 8. Three observed outcomes exhibited by the test overhangs: (a) bonding, (b) curling, and (c) failure to build.to build.

Three observed outcomes exhibited by the test overhangs: (a) bonding, (b) curling, and (c) failure to build.to build. This hints at a gap distance between 0.6 mm and 0.7 mm, where the test overhangs would curl as depicted in Table 2. Curling is still not desirable, but an outcome that is closer to successfully building. Further testing would be needed to investigate the effect of gap distance on the build quality.Table 1: Test Observations with frequency in the test.Test #123Gap (mm) 0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.5 1.0 1.5 2.0 0.4 0.5 0.6 0.7Bonded 3 3 3 3 3 3 3 55 5 4Observation Curled1 1Failed5 5 54508Table 2: Fabrication results for different gap distances.Gap Range (mm)d < ~0.60.6 < d < 0.7d > ~0.7Result:BondingCurlingFailure to buildd = Gap distance4. Method: Flipping and Fixturing (Successful)The next method attempted was one that is termed as “flipping and fixturing,” as shown in Figure 9.

Method: Flipping and Fixturing (Successful)The next method attempted was one that is termed as “flipping and fixturing,” as shown in Figure 9. The general idea is to divide the part with a horizontal cut into two sections that are both void of any problematic overhang. The first step in the fabrication would be to print out the larger of the two sections on a build plate as how normally a part would be printed, and when it is completed the built section would be flipped over, fixtured back onto the build plate where the other section would then be printed on top of it. While a novel idea this potential method came with potential issues and concerns that had to be overcome before the theory could be tested. For testing purposes, the bandpass filters were modified by removal of the holes to facilitate their manufacturability in SLM, and is hereby referred to as the proof-of-concept pieces.Figure 9.

For testing purposes, the bandpass filters were modified by removal of the holes to facilitate their manufacturability in SLM, and is hereby referred to as the proof-of-concept pieces.Figure 9. Illustration of flipping and fixturing methodThe first problem needed to be solved was how the first section of the part will be fixtured to the substrate for the other section to build on top of it. Early ideas included adding a support structure on top of the part similar to the one on the bottom of it so when flipped it would be possible to align the location of the builds. Another idea contemplated was using an adhesive such as soldering or an epoxy. Past the issue of fixturing it to the plate is making sure the fixture is509centered and is exactly even in the z-direction as well. Ideas for this problem include coordinate mapping the build plate and centering on the previous build plate.

Ideas for this problem include coordinate mapping the build plate and centering on the previous build plate. One way to do that would be to create no build zones (parts in Materialise Magics that cannot be built over) with geometry of cut parts on top of where the old builds were as to keep the same alignment and centering. The latest and most executable ideas center around using guiding blocks to center the part. Other ideas have been investigated but not yet experimented such as the “hole” and “screw” methods. The last concern is the quality of bond that will be created after fixturing and flipping. In the study only one method has been attempted, i.e., use of guiding blocks.4.1 Guiding Block ApproachUnder the flipping and fixturing technique, the cut part is built regularly first. Then a second build plate with the guiding blocks outlining the part is built. The cut part is then flipped over and placed into the blocks that act as guiding blocks to center the part.

Then a second build plate with the guiding blocks outlining the part is built. The cut part is then flipped over and placed into the blocks that act as guiding blocks to center the part. Then using the same coordinates from the guiding block build plate, the rest of the build is completed on top.4.2 Part MeasurementsBefore using the guiding block method to fixture the parts to the build plate it was necessary to determine the distance between the blocks and the part that should be allowed. As the parts’ melt pools expand the components slightly it is necessary to put some distance between the part and the block. Measurements were taken using stainless steel calipers corresponding to the dimensions depicted in Figure 10. Each part was measured in five spots for the total width, fourFigure 10. Locations of the measurements taken of the printed part using the flipping and fixturing method.

Each part was measured in five spots for the total width, fourFigure 10. Locations of the measurements taken of the printed part using the flipping and fixturing method. spots the total length, four spots each for the width of the bottom and top, and two spots each for the width of the left and right.The measured results are given in Table 3. The numbers in the table indicate that the widths of the sides are on average 0.0052 mm larger than the CAD dimensions and all the differences were positive, meaning all the sides were wider than the CAD dimensions. The total lengths and widths are on average only 0.0059 mm larger than the CAD dimensions.510Table 3: Measurements of bandpass filters compared to the CAD dimensions.Bandpass Filter ACAD Dimensions (mm) Avg. of Meas.

of Meas. (mm) Differences(mm)Length86.5886.58520.0052Width30.1530.15640.0064Bandpass Filter BLength Width73.46 18.8473.4653 18.84680.0053 0.00684.3 Proof of Concept TrialUsing the guiding block method on the proof of concept part with trial distances between the part and guiding block of 0.005 mm, 0.05mm, and 0.1 mm, the proof of concept part was able to be flipped and fit into each set of guiding blocks. To fit the part into the 0.005 mm guiding block the top of the part (before flipped) needed to be sanded and hammered to fit inside. The 0.1mm guiding block was too loose and no part was attempted in that guiding block.To level the parts vertically, minor sanding was done to ensure there were no errors when printing. An artificial powder bed was made by lowering the elevator and pouring in powder manually then slowly (0.05 mm at a time) raising the elevator until the part was detected visually.

An artificial powder bed was made by lowering the elevator and pouring in powder manually then slowly (0.05 mm at a time) raising the elevator until the part was detected visually. Then ensured the wiper would be able to distribute the powder evenly over the flipped parts. In this trial a build of 2 mm (Figure 11) was printed on top of the cut parts.The trial part was built to completion and the bond evaluated. To evaluate the bond a 5 mm section of the side wall was removed by EDM. Then the removed sample piece was mounted and polished to a 1 µm finish. At this point the sample piece was evaluated under a microscope. Before another microscopic evaluation the sample piece was etched using a 60% HNO3, 40%H2O solution. After microscopic evaluation, Vickers hardness tests were completed.The guiding block method produced the parts with acceptable geometric accuracy and surface finish.

After microscopic evaluation, Vickers hardness tests were completed.The guiding block method produced the parts with acceptable geometric accuracy and surface finish. The alignment was off by a very slight amount and was more precise in the part with 0.005 mm distance between the guiding blocks than the part with 0.05 mm distance.511Figure 11. The proof of concept part after completion using the guiding block method.When evaluating the bond there was no noticeable faults when examining under the microscope at the bond before the etch was completed. After etching the part, the bond area was very distinguishable as the melt pools are visible under microscope and the bond is in the location where the melt pools flip because the part was flipped (Figure 12). Looking at this bond under microscope after etching gave no indication of fault or sign of a bond that was not viable.

Looking at this bond under microscope after etching gave no indication of fault or sign of a bond that was not viable. The Vickers Hardness test gave results indicating the bond was no different from the bonds in the standard SLM part (Figure 13).Figure 12. Both sides of bond area of sample piece observed under microscope (5X) after etching4.4 Guiding Blocks with Actual PartsWith what was deemed as successful results with the proof of concept build the guiding block method was implemented with the actual bandpass filter designs. Considering the512Figure 13. Graphic of sample with Vickers hardness values in the general area they were taken.measurements previously taken as well as the proof of concept results distances of 0.005 mm, 0.01 mm, and .05mm between the part and the blocks were tested to find the best balance of efficiency and dimensional accuracy.

While possible, positioning the edges was increasingly difficult with the smaller distances between the guiding block and part. To combat this problem a new guiding block design was implemented where the edges that line up with edges of the part were filleted as to make it easier to wedge the part into guiding blocks with smaller distances between the part and the blocks. With this newly designed block the bandpass filter designs are able to be fit inside very tight (0.01mm) guiding blocks in approximately three minutes with average human strength hammering of a rubber mallet.With the newly designed guiding blocks the same process as was trialed with the proof of concept was tested. The resulting parts were produced with visible acceptable geometric accuracy and surface finish. The alignment visibly was seemingly perfect and much more precise than the proof of concept. To test this alignment a CMM (Coordinate Measurement System) was used.

The alignment visibly was seemingly perfect and much more precise than the proof of concept. To test this alignment a CMM (Coordinate Measurement System) was used. This machine tracks the x, y, and z coordinates every time the sensor contacts the part being measured. Taking over 600 coordinate points with this machine and using a MATLAB script, a three- dimensional rendering of the part can be made. To judge the surface finish and alignment a best fit line of all the points on each side is made and the coefficient of determination (R2) value is determined. This value calculates the how close the data are to the best fit line from zero to one- hundred percent. The average of these value for the low frequency filter was 0.98, which indicates that the alignment is almost perfect.As with the proof-of-concept trial, the bond was evaluated by using an EDM to remove a sample piece to be mounted and polished to 1 µm finish.

At this point the sample piece was looked at under microscope to be evaluated. Before another microscopic evaluation the sample piece was513etched using a 60% HNO3/40%H2O solution. After microscopic evaluation Vickers hardness tests were completed.When evaluating the bond there were no noticeable faults when looking under the microscope at the bond before the etch was completed. After etching the part, the bond area was very distinguishable as the melt pools are visible under microscope and the bond is in the location where the melt pools flip because the part was flipped. Looking at this bond under microscope after etching gave no indication of fault or sign of a bond that was not viable (Figure 14).Figure 14.

Looking at this bond under microscope after etching gave no indication of fault or sign of a bond that was not viable (Figure 14).Figure 14. Both dides of bond area of sample piece looked at under microscope (5X) after etchingDiffering from the proof of concept where Vickers hardness tests were only completed around the bond area, with the actual part the tests were completed evenly in a 500 by 500 grid throughout the whole cross-sectional area to test if any trends were evident. Creating a color map of the area it is evident there are no trends as the hardness values all fall within 40 units indicating again that bond is viable; see Figure 15.514Figure 15. Vickers Hardness test result5. ConclusionsThe work in this study proves that interdigital bandpass filters can be printed in the SLM process although they are difficult to manufacture due to the overhangs present.

ConclusionsThe work in this study proves that interdigital bandpass filters can be printed in the SLM process although they are difficult to manufacture due to the overhangs present. Generation of optimal support material to aid in the construction of the filters was found to be unsuccessful due to the undesirable part distortion during fabrication and the large surface roughness of the downfacing surfaces. While non-contact supports do appear to have potential in alleviating the distortion encountered, the surface roughness will undoubtedly be present on the downfacing surfaces greatly influencing the performance of the bandpass filters. The solution implemented in this work for overcoming the manufacturing challenges of building overhangs in SLM involved flipping and fixturing, where only a portion of the bandpass filters are printed at a time to bypass building the overhangs.

This technique proved to be dimensionally accurate while still maintaining acceptable bond quality at the interface.AcknowledgementThis work was funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and515the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.References[1]F. Calignano, “Design optimization of supports for overhanging structures in aluminum and titanium alloys by selective laser melting”, Materials & Design, Vol. 64, 2014.[2]Gedda, H., Powell, J., Wahlström, G., Li, W.-B., Engström, H. and Magnusson, C. (2002), “Energy redistribution during CO2 laser cladding”, Journal of Laser Applications, Vol. 14, pp.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAdditive Manufacturing of Metal Functionally Graded Materials: A ReviewYitao Chen*, Frank Liou*Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractFunctionally graded materials (FGMs) have attracted a lot of research interest due to their gradual variation in material properties that result from the non-homogeneous composition or structure. Metal FGMs have been widely researched in recent years, and additive manufacturing has become one of the most important approaches to fabricate metal FGMs. The aim of this paper is to review the research progress in metal FGMs by additive manufacturing. It will first introduce the unique properties and the advantages of FGMs.

The aim of this paper is to review the research progress in metal FGMs by additive manufacturing. It will first introduce the unique properties and the advantages of FGMs. Then, typical recent findings in research and development of two major types of metal additive manufacturing methods, namely laser metal deposition (LMD) and selective laser melting (SLM), for manufacturing different types of metal FGMs will be discussed. Finally, the major technical concerns in additive manufacturing of metal FGMs which are closely related to mechanical properties, and industrial applications of metal FGMs will be covered.Keywords: Additive Manufacturing; Functionally Graded Materials; Mechanical Properties; Laser Metal Deposition; Selective Laser MeltingIntroductionFunctionally graded materials (FGMs) are advanced materials characterized by a gradual change in properties with respect to spatial positions [1].

The idea of FGM originated from composite materials, which combine two or more different constituents and exhibit good material properties [2]. However, sharp transitions between dissimilar constituents often exist in composite materials, which are stress concentrators that can lead to delamination, which is a detrimental material failure [1]. To avoid this disadvantage and eliminate sharp transitions, the concept of FGM was suggested to introduce a gradient zone between dissimilar constituents. Fig. 1 illustrates the difference between composite materials and FGMs. As shown in Fig. 1(a), in composite materials, due to the abrupt switch from material A to material B, the material property function F(x) undergoes a sharp stepwise transition. Sharp transition in material properties at the interface will result in very different responses to the environment at the interface and cause damage.

Sharp transition in material properties at the interface will result in very different responses to the environment at the interface and cause damage. Instead, if we change from material A to material B with a compositional gradient in the middle as Fig. 1(b), then F(x) can be a continuous function, thus relieving the significant difference. FGM was invented by Japanese scientists in the 1980s to design a thermal barrier with a temperature difference as high as 1000 K [5]. Later on, FGMs have been designed to be used in many other areas such as automotive, aerospace engine, implant, etc. [1,6].1215Fig. 1. (a) Composite materials combined with material A and material B: material property F(x) is a step function which changes abruptly at the interface, (b) FGM which gradually varies from material A to material B: material property varies smoothly along the dimension.There are several ways to classify FGMs.

According to the type of grading, FGMs can be classified as stepwise graded FGMs or continuous graded FGMs. FGMs can also be classified into compositionally graded FGMs and structurally graded FGMs [61]. In compositionally graded FGMs, the grading is created by the gradual variation in chemical compositions, while for structurally graded FGMs, the grading comes from graded spatial structure, such as graded porosity and lattice. In Fig. 2, FGM “A” is an example of FGM with graded porosity and here the pores vary gradually in size along the grading direction. FGM “B” is graded in lattice structure and in this case, the thickness of lattice structure increases gradually from the top to the bottom.Conventional methods for manufacturing FGMs include vapor deposition [9], powder metallurgy [7], centrifugal casting [10], etc. Vapor deposition is used for thin-coating FGMs.

Vapor deposition is used for thin-coating FGMs. Due to its high energy consumption and toxic by-products, vapor deposition is not an environmentally friendly approach to produce bulk FGMs [3]. Powder metallurgy and centrifugal casting are used to produce bulk FGMs. Manufacturing FGMs by powder metallurgy involves preparation of the mixture of powder materials according to a predesigned dispersion formulation, stacking, and final sintering [8]. Since the stacking and powder mixing should be prepared before the final sintering process, powder metallurgy can only deal with stepwise grading. Centrifugal casting can realize continuous grading. However, due to the limitation in hardware, only cylindrical sections of FGMs such as tubes can be obtained [11]. The traditional methods mentioned above all need quite a lot of manufacturing tools and assembly processes, which makes the entire procedure time-consuming.

The traditional methods mentioned above all need quite a lot of manufacturing tools and assembly processes, which makes the entire procedure time-consuming. Nowadays, a new approach called additive manufacturing (AM) is attracting a huge amount of interest. AM is defined by ASTM as “the process of joining materials to make objects from 3D model data, usually layer upon layer, as opposed to subtractive manufacturing methodologies [15]”. A 3D computer model of a 3D object can be constructed and sliced into thin layers by computer-1216aided design (CAD) tools. Then the AM system will form each layer under the guidance of the tool path and join all the layers together to make a 3D part.AM is regarded as having major advantages over traditional processing such as its ability to fabricate complex geometries and customized parts, its shorter lead time, less tooling, less assembly required, and overall less waste generated [16].

Due to the important role of metals and their alloys in our everyday life, numerous research works have been done over AM of metal. The two major types of AM for fabricating metallic components are laser powder-fed (LPF) AM and laser powder-bed (LPB) AM [17]. LPF belongs to direct energy deposition (DED) in the seven categories of AM according to ASTM standard, and it is also regarded as laser metal deposition (LMD). In the LPF process, a focused laser beam is applied to create a melt pool on a metal substrate. At the same time, a stream of powder particles is blown into the melt pool by powder feeders. The moving is controlled by CNC for each layer. This process will repeat layer by layer until finally a 3D metal part is achieved. The whole process is carried out in an inert gas atmosphere to avoid metal oxidation. In LPB, a laser beam is used to selectively melt a layer of powder on the powder bed. Thus, LPB is also known as selective laser melting (SLM).

In LPB, a laser beam is used to selectively melt a layer of powder on the powder bed. Thus, LPB is also known as selective laser melting (SLM). The melted powder particles join together and become solid immediately. Once a layer is solidified, the roller spreads a new layer of metal powder to the fabrication zone and the laser melting process will be repeated layer by layer [18]. As this goes on, an as-built 3D part will be generated, along with other un- melted powder particles that can be recycled for future use. That explains why AM generates less waste than conventional subtractive manufacturing.Fig. 2. Structurally graded FGMs with grading in porosity (A) and grading in lattice structure (B).Due to the fast growth of metal AM and metal FGMs, additively manufactured metal FGMs have become a big topic. LMD and SLM are utilized for different types of metal FGMs.

LMD and SLM are utilized for different types of metal FGMs. The LMD process is good at controlling chemical composition by changing the feeding rate for different1217types of powder, but is weak in controlling structures such as pores or lattice. Thus, LMD is a better choice to manufacture FGMs that are graded in chemical composition. For FGMs that are graded in spatial geometry, such as functionally graded lattice structure or scaffold, SLM is the better approach to precisely control the spatial distribution of materials and meet the special structural design requirements [61]. Although there are now numerous review articles on FGMs [1, 3, 4, 8, 12, 13], there is still a lack of comprehensive understanding on additive manufacturing of metal FGMs. This paper presents research work in LMD and SLM to produce metal FGMs with gradings in composition and structure.

This paper presents research work in LMD and SLM to produce metal FGMs with gradings in composition and structure. Critical issues of manufacturing metal FGMs are discussed in detail.FGMs Fabricated by LMDLMD can flexibly control the chemical composition at different deposition positions by gradually varying the ratio of different metal powders. It creates a graded mixing ratio of metal powders from layer to layer. After the fabrication, the graded ratio will result in a graded chemical composition and graded mechanical properties. This technology is especially useful in many industrial areas such as automotive and aerospace industry [14], where multiple types of outstanding mechanical properties are required within one metal component [3]. For instance, a cutting tool needs high hardness and wear resistance at the edge and sufficient toughness at the root.

For instance, a cutting tool needs high hardness and wear resistance at the edge and sufficient toughness at the root. Sometimes two good mechanical properties, such as high hardness and high ductility, are not able to coexist within a part made by homogeneous materials. Therefore, LMD can be used to make new designs of metal components by merging two or more materials together to make metal FGMs according to the special requirements in different portions. This section will review recent important findings in manufacturing metal FGMs by LMD.One of the important aims of designing metal FGMs by LMD is to gradually add a strengthening phase. Titanium alloys, especially Ti-6Al-4V, are widely used because of their low density, high strength-to-weight ratio, and biocompatibility. A decent number of research efforts have focused on the feasibility of manufacturing Ti-based FGMs reinforced by TiC to improve hardness and wear resistance of titanium alloys.

A decent number of research efforts have focused on the feasibility of manufacturing Ti-based FGMs reinforced by TiC to improve hardness and wear resistance of titanium alloys. Early successful work in fabricating Ti/TiC FGM was done by Liu and Dupont [20]. Later the hardness study was performed by [21] and [22]. Zhang et al. [22] deposited functionally graded Ti-6Al-4V/TiC by LMD with stepwise gradient: 10% TiC, 20% TiC and 30% TiC. The variation of hardness along the compositional gradient was studied. Vickers hardness changed from 300 HV1.0 at the Ti-6Al-4V bottom substrate to 600 HV1.0. Zhang et al. [21] fabricated Ti–40vol.%TiC FGM by the LMD process. Different laser parameters were applied in different layers according to the variation of material composition. The hardness of the Ti-40vol.%TiC FGM increased smoothly with the TiC concentration, and no obvious interface was observed. Mahamood et al.

The hardness of the Ti-40vol.%TiC FGM increased smoothly with the TiC concentration, and no obvious interface was observed. Mahamood et al. [23] fabricated Ti-6Al-4V/TiC FGM with composition from 100% Ti-6Al-4V to 50% Ti-6Al-4V/50% TiC. Wear volumes for samples made by fixed parameters and optimized parameters were tested and compared. It was discovered that the optimized parameters can result in better wear resistance. The relationship between tensile properties and TiC volume fraction was studied by Li et al. [24]. The ultimate tensile strength (UTS) of Ti-6Al-4V/TiC FGM with 5% TiC volume fraction is enhanced by about 12.3%. However, when the TiC volume fraction exceeded 5%, the tensile strength of FGM started to decrease. Those representative research works in combining titanium alloy and TiC by LMD are summarized in Table 1.1218Table 1.

Those representative research works in combining titanium alloy and TiC by LMD are summarized in Table 1.1218Table 1. Typical reported research in functionally graded titanium alloy/TiC by LMDComposition/Grading From 100% Ti-6Al-4V to 30% TiC with 10% gradient increment From 100% Ti to 40vol.% TiCDeposition effects/Mechanical properties • Hardness changed from 300 HV1.0 at the Ti-6Al-4V bottom substrate to 600 HV1.0;Crack free. • Hardness HV5 increased smoothly from the substrate along the grading. Hardness doubled at about 20 mm from the substrate; [21]Crack free. • Hardness HV5 increased smoothly from the substrate along the grading. Hardness doubled at about 20 mm from the substrate; [21]The composition of unmelted and resolidified TiC increased with TiC composition.From 100% Ti-6Al-4V to 50% Ti-6Al-4V/50% TiCOptimized functionally graded sample showed the lowest wear volume;From 100% Ti-6Al-4V to 50vol.% TiCMicrohardness was as high as 1200 VHN.

• UTS improved by 12.3% with 5% TiC volume fraction. UTS decreased when TiC volume fraction exceeded 5%; [24]No crack.If two different metals with disparate properties can be combined by LMD, more exotic alloys with novel properties will be generated. Ti-6Al-4V was also used to merge with molybdenum (Mo). Ti-6Al-4V/Mo FGM was deposited by Catherine Schneider-Maunoury [26]. Good metallurgical bonding was found between the deposited wall and the substrate, and also between adjacent deposited layers [26]. Ti/Invar FGM was deposited by Bobbio et al. [27]. Apart from titanium alloys, stainless steel was also investigated to combine with other alloys. The 304L/Inconel 625 FGM, whose composition varied from 100% 304L on one side to 100% Inconel 625 on the other side, was successfully fabricated without a sharp interface by Carroll [28]. Joining dissimilar materials may create different types of intermetallic compounds at the interface that result in properties that are hard to predict.

Joining dissimilar materials may create different types of intermetallic compounds at the interface that result in properties that are hard to predict. Liu et al. [25] studied lightweight Ti/Al graded material for combining excellent properties of both Ti alloy and Al alloy. The composition of Ti/Al varies from pure Ti-6Al-4V to pure AlSi10Mg, with a complex phase change in the middle. Along the composition gradient, the hardness increased and then decreased [25]. Sometimes the aim is to apply functionally graded method to create a high-strength intermetallic phase, such as Ti-Al [19, 29]. In other cases, brittle intermetallic phase, such as Ti-Fe, needs to be avoided which may cause cracks or even fracture such as the joining between Ti-6Al-4V and 304L [30]. Typical combinations between two dissimilar metal alloys which change from 100% of one alloy to 100% of the other alloy are listed in Table 2 in accordance to [25, 26, 27, 28, 30].

Typical combinations between two dissimilar metal alloys which change from 100% of one alloy to 100% of the other alloy are listed in Table 2 in accordance to [25, 26, 27, 28, 30]. From Table 2, it can be found that in most cases, the resultant powder mixing and element composition can be accepted, while cracking often occurs due to unmelted particles or interaction between different types of phases with distinct thermal properties generated during melting and solidification. More details on the brittle intermetallic phase and cracking will be covered in the discussion.1219Table 2.

Summary in combining dissimilar metal alloys by LMDMaterials Ti-6Al-4V & AlSi10MgComposition/Grading Deposition effects/Properties • From 100% Ti- to 100%Element composition gave a nearly linear6Al-4V AlSi10Mgrelationship along the gradient;The maximum hardness reached 619 HV0.1 at the middle section of the deposition; • Zig-zag region in tensile stress-strain curve due to crack;5 gradient regions 25wt.% with incrementFracture surface showed a transition from ductile to brittle along the graded direction.Ti-6Al-4V & MoFrom 100wt.% Ti- 6Al-4V to 100wt.% Mo5 gradient regions 25wt.% with incrementHomogeneous powder mixing; • Satisfactory metallurgical bonding; • Unmelted Mo resulted in the composition of Mo different from expected;Minimum hardness 265 HV at 25wt.% Mo region, maximum hardness 450 HV at 75wt.% Mo region.Ti & InvarFrom 100 vol.% Ti to 100 vol.% InvarActual element composition did not [27]significantly deviate from expected; In the gradient region there was a large fluctuation in element composition due to segregation;Major crack through the gradient due to the presence of secondary phase FeTi and Fe2Ti; • High hardness plateau at the region of 15~33 vol.% and 36~48 vol.% Invar.304L & Inconel 625From 100% 304L to 100% Inconel 625A gradient zone without sharp distinction; • No major deviation in composition; • Cracks of several hundred microns in length were found in a region composed of approximately 79 wt.% SS304L and 21 wt.% Inconel 625.

[28]Ti-6Al-4V & 304LFrom 100 vol.% Ti- 100 6Al-4V to vol.% 304LV was added in 25 vol% incrementsProcess stopped at the region of 50% 304L/50% V;The transition from 25% SS304L/75% V to 50% SS304L/50% V resulted in major cracking at the interface and build failure due to the formation of the σ-FeV phase.Properties other than mechanical properties of FGMs were also studied. For instance, copper can be deposited on steel molds to improve the thermal behavior of steel molds. Since copper and iron are very different in their thermal expansion coefficient, depositing copper directly on steel is very likely to cause failure at the sharp interface. Thus, graded copper was deposited on the H-13 tool steel by Noecker and DuPont [32]. Karnati et al. [33] combined copper and nickel1220Ref.

Thus, graded copper was deposited on the H-13 tool steel by Noecker and DuPont [32]. Karnati et al. [33] combined copper and nickel1220Ref. [25][26]to make Cu/Ni FGM, which possesses both the high thermal conductivity of copper and the strength in high temperature of nickel to create multifunctional alloy in a high-temperature regime. The graded magnetic property was achieved by Heer and Bandyopadhyay [34]. A magnetic compositionally graded bimetallic structure was built via LMD by combining two different types of stainless steel: SS430 and SS316, and the selective magnetic function was then proved to be achievable.For properties in the biomedical area, titanium alloy is a good candidate due to its biocompatibility. People tend to combine titanium alloys with ceramics or other metal elements with special functions to make Ti-6Al-4V a perfect choice for implants. Novel biomechanical properties were achieved through fabricating graded Ti-TiO2 structures by LMD.

Novel biomechanical properties were achieved through fabricating graded Ti-TiO2 structures by LMD. The addition of TiO2 ceramics in titanium not only improved the hardness but also enhanced the surface wettability [35]. Hard and wear-resistant coating consisting of Co-Cr-Mo was deposited on Ti-6Al-4V. All the deposited materials are nontoxic. Living cell density was tested and it was shown that 50% Co- Cr-Mo surface possessed the best combination of biocompatibility and wear resistance [36]. For artificial implants, one important concern is whether the material used for manufacturing implants is safe for human body in the long run. In Ti-6Al-4V, the Al and V elements can cause health issues over a long period of time. Therefore, the nontoxic element molybdenum can be added into Ti-6Al-4V.

In Ti-6Al-4V, the Al and V elements can cause health issues over a long period of time. Therefore, the nontoxic element molybdenum can be added into Ti-6Al-4V. In addition, molybdenum has a lower Young’s modulus than α-Ti, which can also contribute to the decrease of the overall Young’s modulus in order to match with human tissues and reduce the risk of health concerns [26].In the process of LMD, the low powder capture rate is an issue. Using wire to replace powder can be a potential solution. Wang et al. deposited Ti-6Al-4V/TiC FGM by feeding Ti-6Al- 4V wire and TiC powder simultaneously [37]. Copper powder and nickel wire were used by Li et al. to manufacture Cu/Ni FGM [38]. They compared powder-wire deposition with dual powder deposition and proposed that to deposit two different types of metallic materials by the powder- wire method, it is better to use the material with a higher melting point in wire form and the material with a lower melting point in powder form.

Similarly, wire arc additive manufacturing (WAAM) was also tried in fabricating metal FGMs. Shen et al. [39] used a WAAM system to fabricate Fe-FeAl wall with desired mechanical properties. It has similar mechanisms to powder deposition, and it also possesses high deposition efficiency, low cost in material supply, and higher part density.The FGMs mentioned above are all graded in a single direction. Apart from the grading in one direction, other types of gradient can be achieved by adding a new degree of freedom to the AM system. In [40] and [41], metal FGMs with radial grading were fabricated. A functionally graded tube with grading along the tube radius was developed by Durejko et al. [41]. An amended data matrix code (DMC) file was generated to guide the manufacturing system to generate a gradient along the radius rather than the tube axis. Two types of functionally graded 316L/Fe3Al tubes were fabricated.

Two types of functionally graded 316L/Fe3Al tubes were fabricated. Type A varied from pure 316L at the inner surface to pure Fe3Al at the outer surface, while Type B varied in the opposite way. Hofmann et al. [40] developed an AM system with a combination of rotation and linear motion to deposit 304L stainless steel to Invar 36 gradient cladding on an A286 steel rod. The addition of new degrees of freedom can make a more functional LMD system to achieve more flexible deposition.1221FGMs Fabricated by SLMSLM is a powder bed based manufacturing process; thus, SLM is not flexible enough to do real-time controlling of composition. Hence, SLM is not a proper way to fabricate FGMs with graded chemical compositions. Thanks to the extremely high precision of SLM, researchers take this merit in another area of FGMs: functionally graded cellular materials.

Thanks to the extremely high precision of SLM, researchers take this merit in another area of FGMs: functionally graded cellular materials. Functionally graded cellular materials often exist in a cellular form with grading in lattice structure or porosity in 3D space.SLM is a powerful tool to manufacture components incorporating complicated lattice structures that are difficult for traditional machining. Due to the complex 3D geometry, it is too time-consuming to make molds for the casting process, and the assembly is also troublesome. However, SLM can selectively and precisely control the melting and solidification in each layer and finally make differences in solid layers. Apart from periodic cellular structures, which simply repeat a unit cell across the 3D space, cellular structures with gradient have more attractive properties.

Apart from periodic cellular structures, which simply repeat a unit cell across the 3D space, cellular structures with gradient have more attractive properties. They are different from compositionally graded FGMs fabricated by LMD, which usually need fully dense manufacturing to avoid major defects under extreme loading conditions, and cellular materials will be used in certain areas where specific lightweight or hollow structures are required.One important area for research and development on cellular solids is the mechanical response under quasi-static loading and dynamic impact. Fabrication of lightweight lattice structures for energy absorption has also become popular in these years. Previous research has proved the unique mechanical behaviors and capabilities in energy absorption of lattice structures [42]. Recently, researchers have started to add gradients onto these intriguing structures and find more novel mechanical behaviors.

Recently, researchers have started to add gradients onto these intriguing structures and find more novel mechanical behaviors. There have been many cellular structure designs in the past but with no gradient. Thanks to the intrinsic high degree of design freedom and rapid development of manufacturing precision that SLM can provide, we can not only fabricate lattice structures without the constraint of molds but also import gradient to those materials by controlling the volume fraction in different manufacturing layers. By using SLM, we can manufacture graded lattice structures by controlling the 3D geometric design and graded porosity parts by adjusting the building parameters such as laser power, scanning speed, and hatch space, etc.Graded lattice structures are often seen in controlling the 3D geometric design to achieve density gradient. Researchers have reported different types of gradient lattice structure with different unit cells and alloys by SLM. Maskery et al.

Researchers have reported different types of gradient lattice structure with different unit cells and alloys by SLM. Maskery et al. [42] studied mechanical properties of graded Al-Si10-Mg body-centered cubic (BCC) lattice structures manufactured by SLM. They linearly decreased the diameter of struts from the bottom to the top to create a grading. Heat treatment was applied after the SLM process to eliminate the thermal stress and homogenize the microstructure. Both the mechanical behavior of functionally graded lattice and uniform lattice were studied. A uniaxial compression test with a rate of 3 × 10-2 mm/s was applied to the heat treated uniform sample and the heat treated graded sample [42]. Different mechanical behaviors were found after the comparison test. For the compression of the uniform lattice sample, abrupt structure collapse occurred with an angle of 45° when the lattice strain reached 12%.

For the compression of the uniform lattice sample, abrupt structure collapse occurred with an angle of 45° when the lattice strain reached 12%. As the lattice strain kept increasing, the collapse effect became more severe and formed a 45° shear band, which was regarded as abrupt shear failure. By contrast, for the compression of the graded sample, the1222collapse started from the top layer with the thinnest struts to the lower layers with the thickest struts. From the comparison, it is believed that functionally graded lattice has a more reliable and predictable crushing behavior. Thus, it is a more favorable design for structural engineers [42].Choy et al. [44] studied both Ti-6Al-4V cubic lattice and honeycomb lattice in the density gradient made by SLM. The graded structure was also realized by linearly changing the diameter of the struts.

[44] studied both Ti-6Al-4V cubic lattice and honeycomb lattice in the density gradient made by SLM. The graded structure was also realized by linearly changing the diameter of the struts. Comparison between the uniform lattice structure and the graded lattice structure was also made and a similar compressive phenomenon was observed. Additionally, the difference in lattice orientation was covered. The test results showed that FGMs with vertical struts were better at energy absorption [44]. Al-Saedi et al. [43] studied the mechanical properties of uniform and functionally graded structures in the F2BCC unit cell, which consisted of one body-centered cubic (BCC) unit cell and two face-centered cubic (FCC) unit cells. The entire structure was made of Al-12Si and manufactured by SLM. The strut diameter increased gradually from top to bottom.

The entire structure was made of Al-12Si and manufactured by SLM. The strut diameter increased gradually from top to bottom. Experimental work and numerical analysis were performed to understand the novel behavior of additively manufactured functionally graded F2BCC lattice structures under compressive loading. For the functionally graded structure case, the layer-by-layer crushing was successfully modeled by finite element method. However, due to the difficulty in modeling collapse behavior, the finite element result of the uniform case still needs further study [43]. Another type of lattice gradient consists of gradually varying the strut distribution but keeping the strut diameter as a constant value. A typical example work is in [45]. A density graded lattice structure made of Ti-6Al-4V, based on a rhombic dodecahedron unit cell, was fabricated by SLM.

A typical example work is in [45]. A density graded lattice structure made of Ti-6Al-4V, based on a rhombic dodecahedron unit cell, was fabricated by SLM. Both static loading with a strain rate of 0.001/s and dynamic loading with a strain rate of 1000/s were applied during testing, which showed the strain rate sensitivity for the bulk structure.As it has the specific capability to customize manufacturing products, it is worth mentioning the important role of SLM in the biomedical field. Metal materials such as titanium alloy and cobalt chromium are both biocompatible and load-bearing resistant, which attract more attention from researchers who work in designing artificial implants. Due to the large difference in Young’s modulus between solid metals and natural biomaterials, stress shielding may occur during the service period.

Due to the large difference in Young’s modulus between solid metals and natural biomaterials, stress shielding may occur during the service period. Thus, porous or lattice structures are preferred to reduce the overall Young’s modulus of implants and also improve osseointegration in the biological point of view. Moreover, for tissues and organs in human body such as teeth and bones, they are actually natural FGMs with cellular grading [16]. Thus, if metal artificial bones and teeth can be produced with functional grading that can effectively mimic the real biomaterials, the reliability of implants will be raised to a higher level. Controlling the Young’s modulus and yield strength via adjusting the structure gradient of FGM to match the mechanical properties of natural bone structure was mentioned in [49]. The effective elastic modulus of the titanium alloy scaffold was in the range of 0.28~0.59 GPa.

The effective elastic modulus of the titanium alloy scaffold was in the range of 0.28~0.59 GPa. Literature [50] reviewed the state-of-art and applications of additively manufactured FGMs in orthopedic departments. Case studies in designing orthopedic implants by numerical optimization were listed, in which Co-Cr alloy was proposed to produce porosity graded femoral stem by SLM [57, 58]. Both experimental and numerical method proved that the overall stiffness was reduced if the porosity graded design was adopted. Thanks to the rapid development of 3D scanning technology, reverse engineering also contributes to the medical engineering area. In [49], the CT reconstruction was applied to prove a good reproduction for additively manufactured metal implants.1223Another type pf structurally graded structures called porosity graded metal FGMs can be made by controlling building parameters such as laser power, hatch space, etc. As shown in Fig.

As shown in Fig. 3, changing the melt pool size and the distance between melt pools in a graded way in SLM can realize a metal part with gradient porosity. A large porosity can decrease the strength of metals. Then, fabricating porous metal parts with gradient porosity can relieve this problem [46]. Li et al. [46] developed an approach to manufacture steel parts with graded porosity. They verified that the gradient level of porosity can be realized by gradient changed laser scanning speed since when scanning speed increases, the melt pool size will decrease which will lead to controllable pores between melt pools or scanning tracks.Fig. 3.

3. Porosity grading can be controlled by controlling the melt pool size and the distance between adjacent melt pools.Although SLM is mostly used for fabricating FGMs with a grading in lattice structure or porosity, attempts of metal FGMs with other types of gradient by SLM were also made in [48, 54], which may provide new ideas for more diverse design of FGMs. Literature [48] and [54] demonstrated that a graded grain size and texture was achieved by SLM. The graded microstructure was realized by a controllable laser scanning strategy that could maintain both fine grains and coarse grains with clear orientations within one single metal part. Thus, metal parts with flexibly tailorable properties can also be made possible by SLM.

Thus, metal parts with flexibly tailorable properties can also be made possible by SLM. The metal FGM graded in microstructure is not commonly seen; however, it may suggest a novel direction in producing multifunctional metal components.Discussion on Concerns and ChallengesThe LMD process is still a new manufacturing approach, and the mechanisms still need to be understood so that the properties of the produced part made of FGMs can be understood as well [56]. Currently, there is research need for a comprehensive database of additively manufactured metal FGMs [1]. A clear understanding of the material selection and the relationship among chemical composition, phase transition, and mechanical properties is of great importance. As seen in Table 2, when joining dissimilar alloys, unmelted component was often detected which might cause deviation in composition and defects in structure.

As seen in Table 2, when joining dissimilar alloys, unmelted component was often detected which might cause deviation in composition and defects in structure. Optimizing laser processing parameters for depositing two or more different types of powder is critical in order to reach the best joining effect since the key is the interaction between LMD systems and as-deposited FGM parts.1224In particular, cracking is a common issue if the combination is not merged in an appropriate way. For particle reinforced graded alloys, though it is proved that higher particle concentration can result in higher hardness and wear resistance, a limit should be controlled to reduce the amount of unmelted particles and to avoid the crack source [37]. In [35], when the concentration of TiO2 changed from 0 to 100%, cracks were observed. The maximum TiO2 concentration was then reduced to 90% to produce a non-cracked specimen.

In [35], when the concentration of TiO2 changed from 0 to 100%, cracks were observed. The maximum TiO2 concentration was then reduced to 90% to produce a non-cracked specimen. In some cases, intermediate transition route can be applied between two dissimilar materials to avoid specific intermetallic phases that easily cause cracks or even fractures. In [31], a Ti-6Al-4V/V/Cr/Fe/SS316 multi-metallic structure was fabricated to join Ti-6Al-4V and stainless steel 316 in a successful way without generating brittle Fe-Ti phase. In LMD, another important factor that may affect the manufacturing quality is the powder feeding which is a complex process. In this step, due to the difference in powder size and density, the behavior under argon gas flow can be different. There could be powder separation, which will break the predesigned powder ratio and cause a large deviation from the expected metal composition.

There could be powder separation, which will break the predesigned powder ratio and cause a large deviation from the expected metal composition. In [60], the premixed powder flow under the carrier gas was studied numerically and experimentally to provide a guideline for dealing with this manufacturing error. Currently, the majority of FGM parts fabricated by LMD are still test samples. Moving the simple test samples to the real functional components need the comprehensive development of system control. Thus, a global method for LMD path planning should be studied to manufacture functional complex parts rather than single walls [55].SLM is the most commonly used metal powder-bed AM technique. Apart from SLM, other powder-bed based methods can also exhibit their capability in metal FGMs. For generating graded porosity, selective laser sintering (SLS) is a reasonable method as well, especially for FGMs graded in porosity.

For generating graded porosity, selective laser sintering (SLS) is a reasonable method as well, especially for FGMs graded in porosity. The main difference between SLS and SLM is that in SLS, powder particles are only partially melted. Due to the partial melting, there will be more pores generated among powder particles. Erdal et al. [59] applied SLS in manufacturing functionally graded porous polyamide. The application of SLS in metal FGMs is still waiting for more exploration. Electron beam melting (EBM) is also a powder-bed based AM process. The main difference between EBM and SLM is that the energy source used in EBM is an electron beam rather than a laser beam. EBM can be as accurate as SLM, and recently it has also been used for fabricating complex graded lattice structures with high strength and energy absorption [51, 52, 53, 61].In SLM, the key issues mostly lie in material processing, solid mechanics, and design.

For SLM processing, surface roughness and residual stress are major issues. In particular, for cellular graded FGMs made by SLM, a large amount of surface roughness and residual stress can greatly affect and even change the deformation behavior since microscale struts or pores are major constituents. Compared with bulk solid materials, graded lattice structures with fine details can be more sensitive to the size dimensions. Heat treatment processing is necessary for SLM parts. In [44], lattice structures with heat treatment and without heat treatment were compared, and it was found that the compressive response was different between lattice structures with and without heat treatment. It is beneficial to make the step of heat treatment efficient since if the post-treatment takes too long, AM will lose part of its merit of rapid manufacturing. Both surface finish and residual stress can reduce the fatigue life of metal parts.

Both surface finish and residual stress can reduce the fatigue life of metal parts. For graded lattice structures used for long- time service, understanding how much those factors can affect the fatigue behavior is essential.1225Until now, there have been different types of unit cell design for graded lattice structures, such as BCC [42], F2BCC [43], rhombic dodecahedron [45], etc. Most of them are based on experimental methods. With the rapid development of computational methods, numerical modeling is becoming an important tool to reduce material consumption. Topology optimization can be used to find a reasonable unit cell and density distribution design [47]. Topology optimization also has potential applications in designing better support for structurally graded materials with a complicated overhang structure or a multidirectional grading. Finite element modeling is the principal way to predict the mechanical response under different loading conditions.

Finite element modeling is the principal way to predict the mechanical response under different loading conditions. Unlike solid bulk parts, many more possible responses exist in cellular structures, and some behaviors still cannot be accurately simulated. In addition, SLM accuracy is believed to be a factor which results in a gap between experimental results and numerical results [50]. In a word, the development of additively manufactured metal FGMs calls for joint effort among many areas, including development of AM systems, phase transformation, solid mechanics in cellular structures, and numerical modeling.ConclusionFGMs are emerging advanced materials with tailorable material properties that can be used under complex service conditions. Metal FGMs and metal AM are attracting lots of research interest during recent years. Typical research works in different types of metal FGMs manufactured by two major metal AM technologies, namely LMD and SLM are summarized in this paper.

Typical research works in different types of metal FGMs manufactured by two major metal AM technologies, namely LMD and SLM are summarized in this paper. As additively manufactured metal FGMs are still in the early stage, design concerns and challenging technical issues are discussed in order to provide future research directions and arise new topics which can be beneficial to both areas of metal AM and metal FGMs.Reference[1] Rasheedat M. Mahamood et al. Functionally Graded Material: An Overview. Proceedings of the World Congress on Engineering 2012 Vol III WCE 2012, July 4 - 6, 2012, London, U.K.[2] F.C. Campbell. Structural Composite Materials. ASM International[3] Valmik Bhavar et al. A Review on Functionally Gradient Materials (FGMs) and Their Applications. IOP Conf. Series: Materials Science and Engineering 229 (2017) 012021 doi:10.1088/1757-899X/229/1/012021[4] Shailendra Kumar Bohidar, Ritesh Sharma, Prabhat Ranjan Mishra.

Microstructure and mechanical properties of functionally graded TiCp/Ti6Al4V composite fabricated by laser melting deposition. Ceramics International 43 (2017) 16638–16651[25] Yang Liu, Chao Liu, Wensheng Liu, Yunzhu Ma, Cheng Zhang, Qingshan Cai & Bing Liu (2017): Microstructure and properties of Ti/Al lightweight graded material by direct laser deposition, Materials Science and Technology, DOI: 10.1080/02670836.2017.1412042[26] Catherine Schneider-Maunoury et al. Functionally graded Ti6Al4V-Mo alloy manufactured with DED-CLAD® process. Additive Manufacturing 17 (2017) 55–66[27] Lourdes D. Bobbio et al. Additive manufacturing of a functionally graded material from Ti- 6Al-4V to Invar: Experimental characterization and thermodynamic calculations. Acta Materialia 127 (2017) 133-142[28] Beth E. Carroll et al. Functionally graded material of 304L stainless steel and inconel 625 fabricated by directed energy deposition: Characterization and thermodynamic modeling.

ditive Manufacturing of ODS Steels Using Powder Feedstock Atomized with Elemental YttriumSeongun Yang a,b, Donghua Xu a, Dongqing Yana,b, Marc Albert c, and Somayeh Pasebani a,b, 1a School of Mechanical, Industrial, and Manufacturing Engineering, Oregon State University, Corvallis, OR, 97330 b Advanced Technology and Manufacturing Institute (ATAMI), Corvallis, OR 97330 c Advanced Nuclear Technology, Electric Power Research Institute, Charlotte, NC, 28262AbstractThis study investigates the microstructure and mechanical properties of an austenitic ODS steel produced by the Laser Directed Energy Deposition (LDED) process using powder feedstock atomized with elemental yttrium. The Microstructure of the samples was characterized by electron microscopy, and mechanical properties were measured using a tensile test and nanoindentation. Further, the thermal stability of the LDED- produced ODS steels were evaluated.

Further, the thermal stability of the LDED- produced ODS steels were evaluated. As-printed samples showed a cellular structure with Si-Mn-Y-O- enriched nanoparticles that were found to be amorphous. After 100 hours at 1000°C in an argon atmosphere, a partially recrystallized microstructure with a decrease in the number density of Y-O-enriched nanoparticles with crystalline structure was revealed. The as-printed (600 W, 600 mm/min) samples exhibited an ultimate tensile strength of 774 MPa and an elongation at a break of 22%.

The as-printed (600 W, 600 mm/min) samples exhibited an ultimate tensile strength of 774 MPa and an elongation at a break of 22%. A lower ultimate tensile strength of 592 MPa and higher elongation of 42% was measured after 100 hours at 1000°C.Keywords: Laser Directed Energy Deposition, oxide dispersion strengthened alloy, 316L stainless steelIntroductionOxide dispersion strengthened (ODS) alloys have garnered significant attention as potential game-changers in advanced reactor technology due to their fine distribution of nano-sized oxide particles embedded in either a ferritic or austenitic matrix [1]. These nanoparticles effectively impede grain growth and immobilize dislocations, resulting in enhanced strength, even at elevated temperatures [3]. Moreover, the cluster-matrix interfaces of ODS alloys can absorb point defects generated during irradiation, thereby reducing radiation-induced swelling [4].

Moreover, the cluster-matrix interfaces of ODS alloys can absorb point defects generated during irradiation, thereby reducing radiation-induced swelling [4]. Despite the advantageous high-temperature properties, thermal stability, and resistance to neutron radiation damage of ODS steels, they are not yet commercially available or widely utilized. This is primarily due to the complexities and costs associated with conventional ODS manufacturing methods [4]. While alternative manufacturing techniques, such as spark plasma sintering (SPS) [3, 4], have been explored to produce ODS alloys, they offer limited scalability. The poor fusion weldability of ODS steels has also limited their application [5].

The poor fusion weldability of ODS steels has also limited their application [5]. Furthermore, the solid-state friction stir welding process for ODS steels remains suboptimal due to challenges like loss of mechanical properties in the plasticized zone, abnormal grain growth, potential coarsening, and the redistribution of Y-based nanoparticles [6].One of the most widely used AM techniques for creating ODS alloys is laser directed energy deposition (LDED) [1]. In the LDED process, metal wires or powders are deposited onto a substrate to form solid structures. This method is particularly advantageous for producing large components and for repairing and restoring worn or damaged parts. Traditional methods for producing ODS alloys often involve ball milling. However, the use of gas-atomized powder containing elemental yttrium as a feedstock in LDED could potentially eliminate the need for ball milling. Wilms et al.

However, the use of gas-atomized powder containing elemental yttrium as a feedstock in LDED could potentially eliminate the need for ball milling. Wilms et al. [11] utilized metallic powder materials in an Ar-atomized state to explore the feasibility of manufacturing ODS composite materials via the DED process. Conveying nanometer-sized powder materials through conventional powder feeding systems is challenging due to the extensive agglomeration caused by electronic forces [11]. The study also employed mechanical alloying to produce the powder compound using grinding balls and milling containers. As a result, the powder compound consisted of metallic particles surrounded1 Corresponding author (Somayeh Pasebani): e-mail: somayeh.pasebani@oregonstate.edu; phone: +1-541-737-36851478by nanometer-scaled Y2O3 particles. Additionally, the study's findings indicated only minor deformation of the metallic particles during the milling process.

Additionally, the study's findings indicated only minor deformation of the metallic particles during the milling process. Consequently, the study concluded that the powder compound was suitable for the powder feed system [11].Sridharan et al. investigated the internal oxidation of 316L containing Y during the LDED process using gas-atomized powder. They compared results from parts manufactured in both reactive and inert atmospheres [14]. The mechanical properties of the samples were evaluated at room temperature and elevated temperatures, and these were compared with existing literature on wrought ODS alloys [15]. While the ultimate tensile strengths were found to be comparable, the deterioration in yield strength with increasing temperature was more pronounced in the AM-fabricated ODS material. Similarly, Miao et al. [12] determined that 316 steel exhibits both mechanical strength and radiation tolerance.

Similarly, Miao et al. [12] determined that 316 steel exhibits both mechanical strength and radiation tolerance. However, the LDED-manufactured 316L+Y using gas- atomized powder showed greater ductility than all wrought 316L ODS alloys [7].Saeidi et al. [8] and Wang et al. [13] demonstrated that oxide dispersion can be achieved during the LPBF of austenitic stainless steels, resulting in enhanced tensile properties at room temperature. However, the oxides identified were predominantly amorphous SiO2 and exhibited an incoherent and sharp interface with the matrix [8, 10]. Additionally, Ghayoor et al. [9] posited that ODS is a crucial approach for strengthening steel alloys. According to their study, ODS involves dispersing yttrium oxide nanoparticles within a metal matrix. This research also revealed that using selective laser melting (SLM) to fabricate ODS can aid in refining microstructures.

This research also revealed that using selective laser melting (SLM) to fabricate ODS can aid in refining microstructures. Consequently, the study concluded that non-uniform layering and a lack of fusion were responsible for the differences between SLM 304L+5 wt. % and SLM 304L.this study assesses the feasibility of manufacturing a 316L ODS alloy powder made through gas atomization. Its novelty lies in evaluating the mechanical properties of the fabricated samples and using powder- fed laser-directed energy deposition to make a 316L ODS alloy. Therefore, the primary focus of this study is to analyze the printability and mechanical properties of the 316L ODS alloy produced through LDED. Researchers must establish a correlation between processing parameters, and the resulting mechanical properties. The study utilized 316L yttrium oxide ODS powder as the feedstock to generate samples for assessment.

The study utilized 316L yttrium oxide ODS powder as the feedstock to generate samples for assessment. The density, microstructure, and mechanical properties were evaluated to determine the optimal LDED processing parameters for the 316L yttrium oxide ODS. The impact of process parameters and post-process heat treatment on the density, microstructure, and mechanical properties were also examined. Hence, this study is unique as it comprehensively analyzes the manufacturing, mechanical properties, and impact of processes involved in producing a 316L ODS alloy powder.Experimental Procedure1. Preparation of Feedstock and Laser Directed Energy Deposition ProcessGas-atomized 316L stainless steel powder, containing 0.14 wt% yttrium oxide and henceforth referred to as 316LY, was sourced from Sandvik. The chemical composition of the feedstock powder is detailed in Table 1.

The chemical composition of the feedstock powder is detailed in Table 1. The particle size distribution (PSD) of the powder was assessed using a Malvern Analytical particle analyzer (Mastersizer 3000).Table 1. The chemical composition of powder according to Sandvik’s datasheetElement (wt.%)CrNiMoSiMnYOCPSFe316L+2Si+0.14Y17.5010.302.401.931.610.140.070.0190.020.01Bal.For the fabrication of coupons meant for density and microstructure characterization (with dimensions of 20×20 mm), a Meltio M450 powder-wire LDED machine was employed. This machine is equipped with six CW diode lasers, each having a power of 200W and a wavelength of 976 nm. Tensile specimens were also fabricated in compliance with ASTM E8 standards. The build plate used in this study was crafted from stainless steel.

Tensile specimens were also fabricated in compliance with ASTM E8 standards. The build plate used in this study was crafted from stainless steel. The lasers were strategically positioned circumferentially/coaxially at the head, creating a hexagonal pattern on the1479substrate with a working distance of 6 mm. A rotating grooved disk powder feeder was employed to ensure the consistent feeding of the powder feedstock into the melt pool. Argon, boasting a purity level of 99.99%, served as both the shield gas and the powder carrier. To mitigate oxidation, the oxygen concentration within the build chamber was meticulously maintained between 0 and 1%. The LDED parameters, which are elaborated upon in Table 2, were fine-tuned to achieve a relative density exceeding 99% and to prevent the occurrence of micro/macro defects.Table 2.

The LDED parameters, which are elaborated upon in Table 2, were fine-tuned to achieve a relative density exceeding 99% and to prevent the occurrence of micro/macro defects.Table 2. Optimized LDED process parameters used for this studyProcess ParameterValues used in this studyLaser Power (W)600, 700, 800Scanning speed (mm/min)600Powder feed rate (g/min)1Contact tip to work-piece distance (mm)6Laser spot size (mm)0.8Overlap distance (mm)0.56 mmLayer height (mm)0.22. Microstructure characterizationThe density of the LDED-produced 316LY samples was determined using the Archimedes method. Metallographic specimens were prepared following standard grinding and polishing procedures. Subsequently, they were electroetched in a solution comprising 10 wt.% oxalic acid and 90 wt.% deionized water, with an application of 15 V DC for 15 seconds.

Subsequently, they were electroetched in a solution comprising 10 wt.% oxalic acid and 90 wt.% deionized water, with an application of 15 V DC for 15 seconds. The microstructure of the samples was meticulously examined using optical microscopy (Zeiss, Axiotron), scanning electron microscopy (FEI Quanta 3D SEM), and electron dispersive X-ray spectroscopy (EDS).3. Mechanical properties and microhardnessThe microhardness of the LDED 316LY specimens was ascertained using a Leco Microhardness Tester (M-400A) with an applied load of 500 g. The average microhardness value was derived from at least 10 random indentations made on the cross-section perpendicular to the build direction. For evaluating the mechanical properties at various testing temperatures, three dog bone specimens were crafted using wire electrical discharge machining (w-EDM) from rectangular blocks.

For evaluating the mechanical properties at various testing temperatures, three dog bone specimens were crafted using wire electrical discharge machining (w-EDM) from rectangular blocks. These dog bones were designed in compliance with the ASTM E8 standard, featuring a gauge width of 6 mm, a thickness of 1.6 mm, and a gauge length of 32 mm. Notably, the test direction of the dog bones was set perpendicular to the build direction. Tensile tests were conducted on an Instron 5969 machine at room temperature, maintaining a strain rate of 10-3 s-1.4. Thermal stabilityAs-printed tensile specimens, accompanied by a small piece of pure Zr (serving as an oxygen getter), were sealed inside evacuated quartz tubes filled with 99.999% Ar, as depicted in Figure 1. These quartz tubes were subsequently placed in a Carbolite RHF chamber furnace. The samples underwent a heat treatment cycle at 1000°C for 100 hours, with both a heating and cooling rate set at 3.0°C/min.1480Figure 1.

The samples underwent a heat treatment cycle at 1000°C for 100 hours, with both a heating and cooling rate set at 3.0°C/min.1480Figure 1. An evacuated quartz tube with LDED 316Y tensile specimen and Zr (as oxygen absorber) subjected to thermal stability testResults and Discussion1. Powder CharacterizationFigures 2a and 2b show the morphology of the 316LY powder, which consists of spherical particles. Finer satellite particles are observed attached to the larger ones, with a D50 measurement of 73.9 m. The chemical composition of the powder's exterior surface, as depicted in Figure 2b, was analyzed using EDS. The EDS spot analysis results, presented in Figure 2c, confirmed the presence of elemental Y on the surface of the 316L powder.𝜇𝜇Figure 2.

The EDS spot analysis results, presented in Figure 2c, confirmed the presence of elemental Y on the surface of the 316L powder.𝜇𝜇Figure 2. (a) SEM micrograph showing the morphology of 316LY feedstock powder, (b) higher magnification of 316LY powder, the inset shows the fragmentized yttria nanoparticles covering the surface of a 316L particle, and (c) EDS spot analysis on the fragmentized yttria nanoparticles shown in (b)The particle shape, size, surface morphology, and flowability significantly influence the density of parts produced via the LDED process [16]. Consequently, powder flowability is a paramount consideration in the fabrication of components using the LDED technique. The Hausner ratio serves as a metric for powder flowability and is a dependable indicator of cohesion [17].

The Hausner ratio serves as a metric for powder flowability and is a dependable indicator of cohesion [17]. The physical attributes of the powder feedstock, including PSD, apparent density (AD), tap density (TD), and Hausner ratio measurements, are depicted in Figure 3 and detailed in Table 3, respectively. The Hausner ratio is defined as the quotient of tap density and can be determined using the subsequent equation [18]:to apparent density( 𝜌𝜌𝑇𝑇)(1)𝐻𝐻 =𝜌𝜌𝑇𝑇 𝜌𝜌𝐴𝐴Table 3. Physical properties of 316LY feedstock powderPowderApparent Density (g/cm³)Tap density (g/cm³)Hausner ratio316L+2Si+0.3Y3.945 ± 0.0104.436 ± 0.0411.1241481(𝜌𝜌𝐴𝐴)Figure 3. A Malvern PSD histogram of 316LY feedstock powder2. Density of LDED 316LY ODS SamplesRelative density was determined by Archimedes' principle the method. A rule of mixture was applied to the nominal composition of 316LY and used as a bench mark for full density.

A rule of mixture was applied to the nominal composition of 316LY and used as a bench mark for full density. Figure 4 shows the relative density values for the LDED-produced samples. The relative density of as-printed samples at 600 W was found to be 97.75±0.26%, which increased to 98.44±0.64% after thermal stability test. In contrast, the as-printed samples at 700 W showed an initial relative density of 99.02±0.35%. However, after heat treatment, this density reduced to 98.48±0.12%. Interestingly, despite the initial density differences stemming from different laser power settings, both conditions displayed comparable trends after heat treatment.Figure 4.

Interestingly, despite the initial density differences stemming from different laser power settings, both conditions displayed comparable trends after heat treatment.Figure 4. Relative densities of 316L+0.14Y LDED as-printed samplesThe observed increase in porosity after heat treatment at 700 W could be attributed to the higher initial energy input during the additive manufacturing process, potentially leading to more pronounced residual stresses or structural defects that become exacerbated during heat treatment. The comparable final relative densities for both cases after heat treatment imply that the heat treatment process tends to converge towards a specific density, regardless of the initial laser power employed in the Directed Energy Deposition (DED) process.3. Microhardness of LDED 316LY ODS samples1482Figure 5 shows the microhardness values for the as-printed samples (600W, 700W, 800W) and those post thermal stability test.

Microhardness of LDED 316LY ODS samples1482Figure 5 shows the microhardness values for the as-printed samples (600W, 700W, 800W) and those post thermal stability test. As depicted in Figure 2.5, microhardness values consistently decreased from 266±5 HV to 234±19 HV as the laser power increased from 600W to 800W. Following 100 hours at 1000°C, there was a significant drop in microhardness values (approximately 180 HV) for samples at laser powers of 600W and 700W. This decline in microhardness suggests that the heat treatment induced notable alterations in the microstructure, potentially encompassing grain growth and the emergence of voids or defects. This observation aligns with the previously noted decrease in relative density after heat treatment.

This observation aligns with the previously noted decrease in relative density after heat treatment. After the thermal stability test, the 600W (180±12HV) and 700W (179±17 HV) samples displayed comparable hardness values, implying that their microstructural attributes and mechanical behaviors became more aligned, irrespective of their initial as-printed conditions.Figure 5. Microhardness values for 316LY produced at different laser power4. Microstructure CharacterizationThe melt pool boundaries, which were caused by the 90-degree rotational scanning strategy of the LDED process are shown in Figure 6. These boundaries correspond to the interfaces between individual laser beads indicating the laser scanning traces of every other layer, which extend in a perpendicular direction. The depth of the beads was approximately 250 µm, which was equivalent to the lamination thickness used in the LDED process.Figure 6.

The depth of the beads was approximately 250 µm, which was equivalent to the lamination thickness used in the LDED process.Figure 6. Optical micrographs of LDED produced 316LY As-printed: (a, d) 600 W, (b, e) 700 W and (c,f) 800 W1483The optical micrographs shown in Figures 6a and 6c demonstrate the relationship between laser power and the depth and width of the melt pool. As laser power increased, more energy was absorbed by the powder, resulting in higher temperatures and longevity of the molten material leading to a deeper and wider melt pool. Figure 6a-f illustrate the cellular pattern microstructure formed due to the rapid solidification rate in laser-scanned regions, a general characteristic of LDED produced samples.Figure 7.

Figure 6a-f illustrate the cellular pattern microstructure formed due to the rapid solidification rate in laser-scanned regions, a general characteristic of LDED produced samples.Figure 7. Optical micrographs of LDED produced 316LY Thermal aged: (a, c) 600 W and (b, d) 700 WFigure 7 presents optical micrographs that demonstrate changes in the microstructure of 316LY steel resulting from LDED and undergoing a heat treatment at 1000°C for 100 hours. At 600 W and 700 W, the micrographs show similar results. Extended heat treatment causes the melt pool boundaries to disappear. The disappearance of these boundaries indicates significant homogenization of the microstructure, perhaps due to activated diffusive processes during the heat treatment.5. Mechanical behavior1484Figure 8.

Mechanical behavior1484Figure 8. Engineering stress-stain curves of as-printed and thermal aged according to laser power, (a) As-printed LDED 316LY ODS alloy at 700 W, and (b) after 100 hours at 1000℃Figure 8 shows the room temperature stress-strain curve for the as-printed LDED produced 316LY ODS and after 100 hours at 1000oC. The results showed that the 600W-AP specimen exhibited the highest ultimate tensile strength of 774 MPa, with an elongation at a break of 22% at RT. The addition of oxide nanoparticles in the current study improved the ultimate tensile strength through a dispersion mechanism [19]. The sample 600W- TA exhibited a lower ultimate tensile strength of 592 MPa, with an elongation at a break of 42%. The thermally aged specimen 700W-TA showed an ultimate tensile strength of 623 MPa and an elongation of 67%. The tensile results are summarized in Table 4.Table 4.

The thermally aged specimen 700W-TA showed an ultimate tensile strength of 623 MPa and an elongation of 67%. The tensile results are summarized in Table 4.Table 4. Comparison between the mechanical properties of the samples produced in the work with different production processes.YS (MPa)UTS (MPa)Ref.LDED 600W-APLDED 600W-TA611±7.9230±5.2774±3.7592±19.322±0.2 𝜀𝜀𝑟𝑟(%) 42±1.4Current workCurrent workLDED 700W-AP604±13.8760±6.732±2.2Current workLDED 700W-TA250±3.6623±13.267±3.8Current workSLM 316L SS495 - 500610 - 62038 - 48[20]LPBF 304L ODS575±8700±1332±5[23]Wrought 316L SS255 - 31045030[21]As Cast 316L SS26255255[22]The fracture surfaces of the specimens labeled 700W-AP and 700W-TA are depicted in Figure 9. Both samples displayed dimples, indicative of a ductile mode of failure.

Both samples displayed dimples, indicative of a ductile mode of failure. For the 700W-AP specimen, the dimple sizes ranged between 1 and 2.2 µm, while for the 700W-TA, they were between 1 and 2.5 µm. Within these larger dimples, one can observe smaller particles, roughly 0.33 µm across. The presence of these particles inside the dimples hints at possible sites where cracks might initiate, potentially leading to the material's eventual breakdown.(a) As-printed LDED 316LY ODS alloy (700 W)1485(b) LDED 316LY ODS alloy after 100 hours at 1000℃ (700W)Figure 9. SEM micrograph showing the fracture surface of LDED 316LY ODS alloy and dimples, Y-riched nanoparticles with corresponding EDS chemical analysis (pointed by arrow), (a-d) 700W as-printed LDED 316LY ODS alloy, and (e-h) 700W LDED 316LY ODS alloy at 1000℃ for 100 hours.EDS analysis was conducted on these particles, which confirmed the Y-rich composition in Table 5.

This finding could suggest that the oxide nanoparticles contributing to the dispersion-strengthening mechanism are Yttrium-based. The persistence of Y-rich nanoparticles after thermal aging, and their association with potential crack initiation sites, suggests a complex interaction between the nanoparticle dispersion, grain structure, and mechanical properties.Table 5.

Chemical analysis by EDS point analysis performed on the dimples of LDED 316LY 700W as-printed and thermal agedElements (wt.%)FeNiCrMnYSiOAP 142.546.476.690.0035.640.676.27LDED 316LY 700W As-printedAP 259.7111.2712.860.5710.691.651.42AP 348.498.638.960.0025.501.155.81TA 154.048.5511.280.0020.761.501.91TA 238.765.1226.240.0021.430.292.09LDED 316LY 700W Thermal AgedTA 338.025.217.120.8936.820.839.14TA 437.893.7438.271.6213.591.081.43TA 548.307.6611.250.8922.931.045.54Summary and ConclusionThe LDED technique has been explored for its capability to produce ODS alloys, with particular attention to how yttrium oxide nanoparticles, formed through internal oxidation reactions in the chamber, influence the microstructure and strength of 316LY alloy samples; the findings offer valuable understanding into the relationship between these nanoparticles and the microstructure and mechanical properties of the alloy samples.The key findings are summarized below:1486Mo1.721.841.461.966.071.972.382.391) The microhardness of the as-printed specimens was found to be between 234 and 266 HV for laser power of 800W and 600W, respectively.

However, heat treatment caused the hardness values of the LDED fabricated 316LY stainless steel specimen to decrease to 180 HV for a sample that was aged at 1000 °C for 100 hours. 2) The mechanical properties of the as-printed samples were also studied and found to be comparable to the thermally aged condition. For example, the as-printed samples exhibited an ultimate tensile strength of 774 MPa and an elongation at a break of 22% at room temperature for sample 600W-AP. However, the thermal aged samples had a lower ultimate tensile strength of 592 MPa and 623 MPa, along with increased elongations of 42% and 67% at 600 W and 700 W, respectively.3) The fracture surfaces of the 316LY 700W samples, both as-printed and thermally aged, exhibited ductile failure modes, as evidenced by the presence of dimples.

Y-rich nanoparticles were identified within these dimples, suggesting potential crack initiation sites, which indicates their role in the dispersion-strengthening mechanism.Acknowledgment The authors This work was supported by the EPRI. SP acknowledges the financial support from NSF advanced manufacturing program (grant number 1856412). And U.S DOE for financial support. The authors also thank OSU Electron Microscopy Center and the Advanced Technology and Manufacturing Institute (ATAMI) facility staff and director.References: 1. Allen, T., Busby, J., Meyer, M., and Petti, D., Materials challenges for nuclear systems. Materials Today,2010. 13(12): p. 14-23.2. Ghayoor, M., Lee, K., He, Y., Chang, C.-h., Paul, B.K., and Pasebani, S., Selective laser melting of austenitic oxide dispersion strengthened steel: Processing, microstructural evolution, and strengthening mechanisms. Materials Science and Engineering: A, 2020: p. 139532.3.

ditive Manufacturing of Optically Translucent PartsMengqi Yuan1 and David Bourell21School of Mechatronical Engineering, Beijing Institute of Technology, CN 1000812Laboratory for Freeform Fabrication, Mechanical Engineering Department,The University of Texas at Austin, TX 78712AbstractA lithophane is a translucent image created by varying the plate thickness; the image is observed using a back lit light source. Additive manufacturing makes highly complex lithophane fabrication possible. Optical properties of laser sintered polyamide (PA) 12 translucent additive manufactured parts were analyzed and applied to the lithophane fabrication. Several parameters affecting lithophane manufacturing performance are discussed including brightness and contrast versus plate thickness and grayscale level, surface finish quality and manufacturing orientation. Moreover, other thermoplastic semi-crystalline materials were analyzed for LS optically translucent part production.

Moreover, other thermoplastic semi-crystalline materials were analyzed for LS optically translucent part production. Plates and lithophanes were built using a different AM platform: stereolithography (SL) with Somos® ProtoGen™ O- XT 18420 white resin. Different optical properties and lithophane performance were observed and compared with PA 12 parts.It was found laser sintered polyamide 12 optical properties varied with light wavelength and reached maximum transmission under green light. When building in the XY plane, thin layer thickness and large maximum plate thickness led to higher contrast and gray scale level. Lithophane quality was largely improved when fabricated in the ZX/ZY plane orientation.

Lithophane quality was largely improved when fabricated in the ZX/ZY plane orientation. Lithophanes made from SL were analyzed but showed lower contrast due to the optical property characteristics of the white resin.IntroductionLaser sintering (LS) is an additive manufacturing technology that uses a laser to fuse polymer powder into a mass that has a desired three-dimensional shape. The laser heating results in melting and powder local consolidation. Polyamide 12 is a thermoplastic material that is widely used in laser sintering. It has high elongation, good specific strength and melts at a temperature around 180°C [1]. A lithophane is a translucent image created by varying the plate thickness; the image is observed using a back lit light source. It is a design in gray tones [2,3].

A lithophane is a translucent image created by varying the plate thickness; the image is observed using a back lit light source. It is a design in gray tones [2,3]. It presents a three-dimensional image that completely different from two dimensional engravings and daguerreotypes that are ‘flat’ [2-4].The optical properties of the material are essential for lithophane application as this is critical to define the brightness and contrast. The fraction of incident light transmitted by an optical material depends on losses due to absorption and back reflection [5].

The fraction of incident light transmitted by an optical material depends on losses due to absorption and back reflection [5]. The transmittance largely varies with wavelength and materials, which is the ratio of the light transmitted versus the incident light intensity according to the Beer- Lambert law:T = (1-R)2 e-αx (1)where T is the fraction of light transmitted, R is the fraction of light reflected, α is the absorption coefficient which depends on the absorber chemical composition and the light wavelength, and x is the path length (i.e., plate thickness for lithophanes).

The total light transmitted then may be written by applying the Taylor’s series expansion:T = (cid:4666)(cid:2869)(cid:2879)(cid:2902)(cid:4667)(cid:3118)(cid:2915)(cid:3127)(cid:3209)(cid:3182)(cid:2869)(cid:2879)(cid:3019)(cid:3118)(cid:3032)(cid:3127)(cid:3118)(cid:3209)(cid:3182) (2)For a given absorption coefficient and refractive index, transmittance can be solved usingthis modified Beer-Lambert Law.A grayscale digital image is an image in which the grayscale levelof each pixel has a single value, which means it carries only intensity information [6]. Brightness is a visual feature which may be described as the average grayscale level of an image. Since the range of grayscales is 0-255, the image will tend to wash out if large numbers of pixels are changed to become either 255 (white washout) or 0 (black washout).

Since the range of grayscales is 0-255, the image will tend to wash out if large numbers of pixels are changed to become either 255 (white washout) or 0 (black washout). Contrast is an indication of the change in luminescence of an image relative to its surroundings.The software utility, Bmp2Cnc, was used to convert the grayscale jpg/bmp digital photo to a three-dimensional CAD file. It is based on a relation that linearly translates each pixel grayscale level to a respective height. It reads the file in bitmap format, generally 8 bits. Z-axis depth is changeable based on the manufacturing requirements.ISO/ASTM 52921: 2013 (E) [7] establishes the standard coordinate systems and test methodologies for additive manufacturing processes.

The part initial build orientation is described by listing which axis is parallel to the longest overall dimension of the bounding box first, followed by the axis which is parallel to the second longest overall dimension of the bounding box second, followed by the axis which is parallel to the third longest overall dimension of the bounding box.Two LS building orientations were used in this research: a ‘flat’ build in the XY plane, and a ‘standing’ build on edge in the ZX/ZY plane. ‘Standing’ manufacturing orientation results in better surface quality and enlarged contrast but requires more time and material consumption.

‘Standing’ manufacturing orientation results in better surface quality and enlarged contrast but requires more time and material consumption. ‘Flat’ building leads to lower contrast due to the layer limitation but is accomplished with less time and material.In this paper, work has been done to analysis the optical properties of laser sintered polyamide 12 [8], and the manufacturing issues related to lithophane performance [9,10] are overviewed and discussed.Optical PropertiesThe optical properties of optically translucent laser sintered polyamide 12 were analyzed by conducting monochromatic light transmission experiments on laser sintered polyamide 12 blank plates.

Blank plates were manufactured on a DTM® 2500 SinterStation, with dimensions 50.8 mm by 50.8 mm and with thicknesses of 0.38, 0.64, 0.83, 0.92, 1.30, 2.33, 4.74 mm,corresponding roughly to 4, 6, 8, 9, 13, 23 and 48 layers of powder during laser sintering with 0.102 mm per layer thickness (system error 5% on z-direction is included). A Unico 1200 110V/50Hz S-1201 spectrophotometer was applied to investigate the light transmittances. Measurement is made from 400 nm to 800 nm in 20 nm increments.The effect of the wavelength on laser sintered polyamide 12 transmittance is shown in Fig. 1. Transmittance first increased linearly from 400 nm to the peak of 540 nm (25- 87%) and dropped to the lowest at 580 nm, then remained almost constant from 600 nm to 800 nm. At a constant wavelength, transmittance decreased monotonically with increasing plate thickness.

At a constant wavelength, transmittance decreased monotonically with increasing plate thickness. It is proposed that the pi-pi* transition is responsible for the transmittance increment from 400- 540 nm, while the n-pi* transition explains the transmittance decreasing from 540- 580 nm [8].The influence of plate thickness on laser sintered polyamide 12 transmittance is shown in Fig. 2. The transmittance of PA 12 plates with a common soft white bulb is presented as well. The white light effect was also calculated based on the theory that white light is a combination of all lights and there is no clear boundary between one color to another [11]. The calculation results match well with the soft white bulb experimental results. All plots show that transmittance decreases with increasing plate thickness.

The calculation results match well with the soft white bulb experimental results. All plots show that transmittance decreases with increasing plate thickness. Green light performs the best, showing the largest sensitivity to thickness from 85% to 20% with increasing plate thickness; the soft white bulb transmitted 20% less light compared to the green light, followed by blue, red and yellow light. The transmittance results are verified in Fig. 3 which shows large visual differences on the actual lithophanes.To conclude, transmittance of laser sintered polyamide 12 decreases with increasing plate thickness. Second, transmittance reaches the highest levels under monochromatic green light.

Second, transmittance reaches the highest levels under monochromatic green light. It reaches the lowest levels under yellow light.Figure 1: Effect of spectrophotometer wavelength on transmittance of laser sintered polyamide 12 plates with different thicknesses.Figure 2: Effect of laser sintered polyamide 12 plate thickness different transmittance monochromatic lights and white light. Also included is light based a calculated dependence for white averaging red, green, and blue measurements. All calculated curves are based on Equation 6 with absorption coefficients obtained from the experimental measurements.onunderFig 3: Gray-scale processed laser sintered polyamide 12 3.81mm thickness lithophanes under different monochromatic light: green (left), yellow (center).

On the right is the initial digital image used to create the lithophanes.Parameters Affecting AM Lithophane PerformanceLayer and Plate Thickness EffectTo analyze the layer thickness and maximum plate thickness effects on laser sintered polyamide 12 lithophane performance, lithophane STL files were generated using a reproduction of Raphael’s painting, Agnolo Doni. The lithophanes were built on a SinterStation HiQ using different layer thickness, either 0.076 mm (15 W laser power) or 0.102 mm (24 W laser power) and different maximum plate thickness (1,3 and 5 mm). Lithophane images were captured using a SONY DSC- W55 camera. A Picker 240050 X-ray view/ light box was used in this research, which produced non-adjustable visually uniform white light over the view area.Fig.4 shows the lithophanes built in the XY plane with maximum plate thicknesses of 1, 3 and 5 mm, which corresponds to 10, 30, and 50 layers, respectively.

The top group shows higher contrast compared to the bottom group.Boundaries and features on the face, tie and hands were lucid. With a constant maximum plate thickness, thinner layer thickness results in more grayscale levels on the lithophane since there are more distinct layers in the part.Moreover, large maximum plate thickness leads to improved contrast and quality since the grayscale levels were enriched. However, if the lithophanes become too thick, problems arise with the viewer-facing surface topography becoming too distracting. Figure 5 illustrates this for four lithophanes, with maximum plate thicknesses of 3, 5, 7, and 9 mm with a front lit white light source and black background. The topography for the lithophanes built with 7 and 9 mm plate thickness become distracting, and the light transmission is very low.

The topography for the lithophanes built with 7 and 9 mm plate thickness become distracting, and the light transmission is very low. There is then a limit on the maximum thickness of the lithophane based on several considerations [10]: for laser sintered polyamide 12 lithophane fabrication, the optimal maximum plate thickness ranged between 3.8 mm and 5 mm.Figure 4: Polyamide 12 lithophanes laser sintered with (a) 0.076 mm layer thickness; (b) 0.102 mm of layer thickness. Maximum plate thickness of each pair was 1, 3 and 5 mm from left to the right. The light source was a white light box.Figure 5(right): Lithophanes presented with a front lit white light source fabricated in the XY plane with the maximum plate thicknesses of (a) 3 mm and (b) 5 mm; fabricated in the ZX plane with the maximum plate thicknesses of (c) 7 mm and (d) 9 mm.

All lithophanes were fabricated with a layer thickness of 0.102 mm.Manufacturing OrientationTo analysis the manufacturing orientation effect on the lithophane surface quality, lithophanes generated from the reproduction of Raphael’s painting, Agnolo Doni, were first manufactured in the XY plane and then in the YZ plane and the ZX plane, with maximum plate thickness of 3.8 mm. Fig.6 shows better contrast and finer surface quality on lithophanes built on the ZX an YZ planes compared to the one built on the XY plane. The background and figure jackets on the first two were more identifiable. Lithophanes built in the ZX plane and XY plane were examined using stereomicroscopy with 2.5 X magnification on the portrait tie and right middle cloud area shown in Fig.7. The lithophane built in the ZX plane presents a much smoother surface and less contouring than the one built in the XY plane. Individual layers were observed on the lithophane built in the XY plane, with 38 layers in total.

Individual layers were observed on the lithophane built in the XY plane, with 38 layers in total. The layers appear like stairs, and each layer defines a quantized grayscale level. The grayscale level of the lithophanes built in the ZX plane was defined by the laser beam motion, so a more continuous level variation was produced.Figure 6: Polyamide 12 lithophanes laser sintered with 0.102 mm of layer thickness built in the ZX, YZ, and XY planes from left to the right. The size of lithophanes was 127 mm by 76 mm, with maximum plate thickness equal to 3.8 mm. Lithophane photos were captured using white light box illumination.Figure 7: Stereomicroscopy photos of the lithophanes built in the XY orientation (left) and ZX planes (right). The focuses are on the portrait tie (top sequence) and cloud on the middle right of the image (bottom sequence).It is reasonable to achieve different feature resolutions depending on the manufacturing orientations.

Small features in the ZY plane were built using continuous laser beam line scan, while the ones built in the XZ plane had a discontinuous/short line scan on the irregular surface figures. Small features were more achievable with long scan lines, since it was continuous. Long line scanning on rectangular features also led to greater precision on positioning. Thus, small features (diameter or width of 0.13 mm, for example) could be produced for the ZY orientation. It is concluded that for LS lithophane generation, manufacturing “on edge” in, for example, the ZX or YZ plane results in better surface quality on the lithophane faces, especially for the thin parts, but the thinnest feature size is 0.57 mm due to the laser beam diameter limitation.

Manufacturing in the XY plane, the thinnest part is roughly one layer, but parts with only one layer thickness are extremely friable.Surface Finish QualityTo analyze the effect of surface finish on lithophane performance, five blank plates were built with dimensions, 50.8 by 50.8 by 12.7 mm. The polishing experiment was conducted at Harvest Technologies, Belton, Texas. Blanks were first infiltrated with cyanoacrylate to obtain a brittle surface to polish, since the polymer tended otherwise to deform and create strands or scratches when sanded. As the cyanoacrylate soaked into any surface porosity of the part, a change in part transmittance was foreseeable.An initial experiment was designed to analyze the cyanoacrylate effect on the laser sintered PA 12 transmittance. A portion of a flat specimen was infiltrated on one surface, and a second piece was infiltrated on both surfaces.

A portion of a flat specimen was infiltrated on one surface, and a second piece was infiltrated on both surfaces. To investigate the polishing effect, a third piece was infiltrated and polished on one side, while a fourth piece was infiltrated and polished on both sides. A fifth piece was neither infiltrated nor polished as a baseline specimen. The optical transmittance change of all four post-processed plates and the non post-processed plate was measured using the spectrophotometer with varying monochromatic wavelengths. Thickness was measured, and its effect on transmittance was examined. The polishing effect was also measured visually by viewing the polished plates using the white light box. Photos were taken with the polished surface facing and away from the light source.Fig. 8 shows the surface roughness results for the non-post-processed (original), cyanoacrylate infiltrated, and polished surfaces. The surface roughness decreased to a minimum after polishing.

8 shows the surface roughness results for the non-post-processed (original), cyanoacrylate infiltrated, and polished surfaces. The surface roughness decreased to a minimum after polishing. However, the transmittance variations for five blanks at the same wavelength were less than 5%. The non-post-processed plate had a rougher surface finish compared to both face polished plates. Small dimples were present on the non-post-processed plate. The post processes (infiltration, infiltration then polish) only dealt with the plate surface with no significant effect observed on the laser sintered polyamide 12 optical properties. It is unlikely to have a smooth surface finish for the laser sintered polyamide 12 lithophane because polishing is difficult to achieve on curved surfaces.Improvement of Lithophane QualityThe definition of a ‘good’ lithophane surface quality may be proposed based on the above experiments and observations, which are summarized as:1.

Light transmittance should be high (> 70%) at the thin areas and low (< 15%) at the thick sections.Figure 8: Surface roughness comparison among the non post-processed, cyanoacrylate infiltrated, and polished surfaces. Error bars representing standard deviation of 5% are shown. The surface roughness is quantified by the vertical deviations of a real surface from its ideal form. Ra is the average of the absolute vertical deviations, Rq is the root mean squared value of the vertical deviations, and Rz is the sum of the highest and lowest vertical deviations within a sample length.2. The number of grayscale levels should be large, which leads to more continuously varied thickness.lithophane maximum plate thickness should fall within a range, below which would be too few grayscale levels, and beyond which the lithophane topography would be distracting.3. The4.

The4. The resolution ofthelithophaneshould be relatively high.In summary, it is proposed that high quality is associated with the lithophane appearance features (brightness, contrast and resolution) under the backlit light source condition that are identical to the original digital image.Based on the definition, the material selection criteria for producing a ‘good’ lithophane was carried out, from which the optimal material was found to have an absorption coefficient of 0.5/mm [10]. A group of laser sinterable candidate materials was selected. The primary criteria included the polymer powder type (semi-crystalline or crystalline, thermoplastic) and melting point (160-200°C). The candidate polymers included EBA, EVOH, PA, PCL, PE, PP and PVDF. Polyamide 12 best fits the selection criteria for lithophane laser sintering, followed by EBA, EEA, and PVDF.

The candidate polymers included EBA, EVOH, PA, PCL, PE, PP and PVDF. Polyamide 12 best fits the selection criteria for lithophane laser sintering, followed by EBA, EEA, and PVDF. It is proposed that nanocomposites might lead to improved properties by adding optical translucent nanoparticles into the base thermoplastic matrix [10].Stereolithography (SL)Stereolithography (SL) was performed as a comparison to the LS manufacturing method. Blank plates and lithophanes were built at Harvest Technologies, Inc. DSM Somos® ProtoGen™O-XT18420 was used in this study. It is a white, liquid, epoxy resin photopolymer. The effect of the light wavelength on SLA Somos® 18420 white resin transmittance is shown in Figure 10, which has a very different transmittance trend. Comparing the plates made with LS and SL, the LS plate (0.38 mm thick) had higher light transmission than the SL plate (0.34 mm thick) from 460 to 800 nm light wavelength.

Comparing the plates made with LS and SL, the LS plate (0.38 mm thick) had higher light transmission than the SL plate (0.34 mm thick) from 460 to 800 nm light wavelength. The 4.71 mm LS blank had consistantly lower transmittance than the 5.11 mm SL blank for all visible light wavelengths. These light transmission differences were shown on the lithophane performances as well.Two lithophanes were built in the XY and ZY planes at room temperature. All parts were immersed in 99% ethanol for 3 hours after build completion; support structures were then removed manually. Images in Fig. 11a and 11b show lithophanes built in the XY and YZ planes using SL, with a maximum plate thickness of 5 mm with layer thickness of 0.102 mm. The image on the far right is a comparison lithophane built in the YZ plane using LS. Lithophanes built with SL show clear detail and no contours on the background, due to the smaller laser beam diameter (0.2 mm) compared to that for LS (0.5 mm).

Lithophanes built with SL show clear detail and no contours on the background, due to the smaller laser beam diameter (0.2 mm) compared to that for LS (0.5 mm). Smaller laser beam diameter leads to more precise scan positioning and thus better lithophane appearance. The lithophane boundaries, figure jacket and hair were darker on the laser sintered lithophane due to the light transmission differences reported in Fig. 10.Figure 12 shows magnified images of lithophanes made using SL built flat and vertically, compared to a vertical LS plate. The vertically built SL plate has more support structures to remove than the SL plate built flat, especially for the SL YZ sample. Also the surface qualities of the lithophane made using SL and LS appeared different, due to the differences of the raw material state (Somos® 18420 is liquid and PA 12 is solid).Since the surface qualities were similar, it is optimal to build the lithophane in the XY plane than the ZY plane.

The lithophane made with SL using DSM Somos®ProtoGen™O- XT18420 white epoxy resin showed good and similar resolution when built in the XY or YZ plane. Violet light (wavelength 400 nm) transmitted best among all visible lights. However, the lithophane contrast was lower than that of the LS lithophane due to low light transmission in the resin.spectrophotometer Figure 10: Effect of wavelength on transmittance of SL resin plates The withthreedifferentthicknesses.Figure 11: SL lithophanes built (a) in the XY plane; (b) built in the ZY plane; and (c) LS lithophane built in the ZY plane. The maximum plate thickness for each was 5 mm.

The maximum plate thickness for each was 5 mm. Lithophanetransmittances of laser sintered 0.38 mm and 4.71 mm thick polyamide 12 plates (dashed lines) are shown for comparison.photos were captured using a white light box.Figure 12: Stereo-micrographs of the lithophane made using SL and LS focusing on portrait hands.Summary and ConclusionsThe objective of this paper is to overview the optical properties, manufacturing issues, and AM manufacturing method relative to laser sintered polyamide 12 lithophane production.

First, the optical properties of laser sintered polyamide 12 were analyzed: the light transmission decreases with increasing plate thickness; also, the transmittance reaches the highest levels under monochromatic green light, which was verified by observing lithophane performance under different monochromatic LED lights.Second, the manufacturing issues for improving the laser sintered polyamide optically translucent part performance were investigated: thinner layers and relatively thicker plates result in finer lithophane performance. However, if the plate becomes too thick, it leads to poor light transmission, and the topography would be observed with front lit light. Sintering a relatively thick lithophane (<5 mm) in the ZX/ZY plane significantly increases the contrast and resolution compared to sintering in the XY plane.

Sintering a relatively thick lithophane (<5 mm) in the ZX/ZY plane significantly increases the contrast and resolution compared to sintering in the XY plane. The thinnest feature thickness possible in the SinterStation HiQ is in the XY plane 0.13 mm with 0.102 mm layer thickness, and it is 0.57 mm when manufacturing in the ZX/ZY plane with the laser diameter of 0.50 mm.The surface finish has no effect on the lithophane quality, since the polishing effect may only be applied on the polished flat surface. Laser sinterable material selection criteria were carried out, from which the optimal material was found to have an absorption coefficient of 0.5/mm. PA 12 was the best semi-crystalline material after comparison with a list of other semicrystalline and crystalline candidate materials. The use of nanocomposites for optimal lithophane performance and more precise manufacturing processing may improve lithophane quality.Stereolithography was analyzed using Somos® 18420 white resin.

The use of nanocomposites for optimal lithophane performance and more precise manufacturing processing may improve lithophane quality.Stereolithography was analyzed using Somos® 18420 white resin. Transmittance reached the highest levels for violet light. Improved resolution and inferior contrast were observed compared to laser sintered lithophanes.Overall, this work investigated the optical properties and manufacturing orientation effect on laser sintered polyamide 12 optically translucent parts, and provided solutions forphotographic reproduced lithophane surface quality improvement. The work is the first discussion on the factors that affect the fabricated part dimension accuracy, as well as the initial description of the manufacturing orientation effects on surface quality for lithophane production to guide do-it-yourselfers and end users of 3D printers in reproducing their own images.

It introduces lithophane additive manufacturing to a wide audience, and provides a durable, high quality, long-lasting photographic record. The study provides analysis of laser sintered lithophane production and serves as a basis for future commercialization art and end- user parts.AcknowledgmentsThe authors would like to thank Billy Herring and Jeff Rodocker from Harvest Technologies for their help and guidance in the SL processing.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperADDITIVE MANUFACTURING OF PERSON SPECIFIC DIABETIC FOOT INSOLES WITH ADJUSTABLE CUSHIONING PROPERTIES USING TPMS LATTICE STRUCTURESF. R. Claybrook*, D. J. Southee*, and M. I. Mohammed*School of Design and Creative Arts. Loughborough University, Loughborough, LE11 3TUAbstractComplications associated with diabetes are numerous, including foot problems which in extreme cases can lead to amputations. Current management involves the use of foam diabetic foot insoles (DFI) to provide cushioning, however load bearing capacity is limited, and designs often do not provide a comfortable or efficacious fit. This study aspires to resolve problems using digital fabrication workflows.

This study aspires to resolve problems using digital fabrication workflows. The exploration of potential 3D scanning of anatomical data, parametric modelling, and additive manufacturing was created for a patient specific DFI. This demonstrated that patient scanning data provides means to create a custom fitting insole template, improving overall fit. Demonstrating the use of triply periodic minimal surface (TPMS) structures, fabricated in Thermoplastic Polyurethane (TPU), as cushioning structures, whereby unique lattice designs allow regionally tailored mechanical loading properties of the insole concept. The final insole realises a superior alternative to tradition DFI.1. IntroductionThe complexity of anatomical representation of the body is one that has been researched to improve clinical processing and burdens on healthcare providers (Mohammed et al., 2016). Diabetic patients increase complications in orthotics with the presence of Peripheral Neuropathy.

Diabetic patients increase complications in orthotics with the presence of Peripheral Neuropathy. Current diabetic insoles are established by the International Working Group on Diabetic Foot 2015 guidelines dictating the requirement of custom-made therapeutic footwear to relieve pressure <30% and the importance of preventive techniques, though the research is sparse (Bus et al., 2015). The current rehabilitation techniques involve patients to daily self- examine and wear properly fitting footwear (Schaper, et al 2016). Yet these intervention techniques are perturbed by cost, time, dexterity, and health impairments, resulting in low conformity with upkeep (Salles & Gyi, 2013; Van Netten et al., 2013).Prosthetic based rehabilitation techniques using insoles requires laborious and intensive production process.

The methodology is costly with a dedicated infrastructure for negative impressions, plaster infill, and vacuum formation for therapeutic footwear (Farhan et al., 2021), limiting manufacturing to highly skilled professionals (Davia-Aracil et al., 2018). The plaster is reusable with the breakdown of the substance for repetitive use, but this destruction of the mould means the anatomical data of a patient is not kept for repeat prescriptions and can increase wait times. The workflow of traditional fabrication techniques has been summarised by Mohammed et al., 2018. The complex process of plaster cast can be defective in the fitting and sculpting process and forming can cause discomfort; even consequential movements from the patients can differ the accuracy of this application process.

Too much variability in successful casting makes the process too dependent on the rectification process, indicating a better need for a clinical workflow (Farhan et al., 2021).1717The recent development of computer aided design (CAD) allows the advancement of 3D replicated scanning and manufacturing to have a higher customisation level than foams on the market by customising to microporous architecture (Mohammed & Gibson, 2018; Sala et al., 2022). Foot scans can be based on geometric point clouds to retrieve the best developed fit (Rout et al., 2010) whilst keeping a non-invasive approach to be easily adopted into clinical practice. Current overlapping technologies involve gypsum plaster and 3D scanning; gypsums white matte appearance is ideal and creates and an excellent skin replication (Paterson, 2013). The use of a replicated limb allows more leverage for scanning, lowering obstruction and mobility to capture anatomical data.

The use of a replicated limb allows more leverage for scanning, lowering obstruction and mobility to capture anatomical data. With scanning technologies becoming cheaper and more accessible to the public (Haleem & Javaid, 2019), the medical application of this technology can potentially be complementary to the field. Research can take advantage of exploring a full augmentation for a complete digital fabrication process without the need for physical casting. An adaptation of Paterson, 2013 for digital orthoses application for gypsum casting has been adapted for full augmentation in Figure 1.

An adaptation of Paterson, 2013 for digital orthoses application for gypsum casting has been adapted for full augmentation in Figure 1. This will be explored for the study.The user chooses between orthoses templates.The patient’s anthropometric scan and other information is collected.Key anatomical landmark points are located, indicated, or imported.If more data is requiredAesthetic design and additional details added.Point cloud of scan is used to generate orthoses outline and modified to create customAlignment and scale of scan is corrected and realigned if needed.Figure 1 - Adaptation of Paterson, 2013 3D CAD flow for 3D scanning orthoses applicationThe use of fuse fabrication filament (FFF) modelling allows improved technologies and the use of lower shore hardness and malleable materials, such as Thermoplastic Polyurethane (TPU) (Harris et al., 2019).

The small, compact, and sterile environment created by FFF is favourable to the medical sector, providing a quicker and higher customisable option of tooling, whereas additive manufacturing. Traditional manufacturing manufacturing (AM) is advantageous as it is tool free manufacturing and can reproduce high levels of complexity, such as lattice structures, ideal for tailored specific energy absorption (Bates, et al., 2016; Claybrook, et al., 2022). The high elasticities and energy absorptions characterised by TPU creates a polymer suitable for applications to control mechanical deformations (Beloshenko et al., 2021). The use of TPU lattices can further cushioningislimited by1718applications, alongside 3D scanning creating a heavily customised insole approach, both mechanically and physically.Lattice structures provide a high strength-to-weight ratio but there has been limited research with triply periodic minimal surface (TPMS) structures (Beloshenko et al., 2021).

The development of the TPMS network and control of porosity can lead to high level customisation that are attractive to the orthotic field. The complex geometry of TPMS structures creates a porous structural unit capable of the customisation needed for diabetic insoles (Ma et al., 2019). TPMS structure, TPU, and FFF are currently limited in studies (Claybrook et al., 2022; Holmes et al., 2022; Sala et al., 2022) and shows an opportunity in this field.In this study we aim to investigate the limited guidelines for orthoses scanning techniques. Our techniques will propose less material waste, time, and labour than required for the traditional insole application techniques. The main objective is to understand the capabilities of the 3D scanning and the digitised process to accurately mimic the plantar surface foot for highly accurate custom-made insole.

The main objective is to understand the capabilities of the 3D scanning and the digitised process to accurately mimic the plantar surface foot for highly accurate custom-made insole. A specification will be developed for the requirements of a successful 3D scan for orthotic application, similarly the preliminary testing of an initial TPMS structure at insole height to understand the mechanical properties achievable.2. Experimental Methodology 2.1.Topography Scanning MappingWe initially assess the capabilities of professional 3D scanners, the Artec Leo, to identify limitations and boundaries of 3D scanning techniques. There is no widespread common practice for foot 3D scanning foot orthoses, therefore reported evidence from literature indicating high reliability of anatomical capturing foot data was followed (Mohammed & Elmo, 2020).

To the authors knowledge, there is only one systematic review comparing reliability and accuracy of 3D scanning with traditional techniques for therapeutic footwear. Farhan et al., 2021 reported there was no evidence for accurate fabrication of foot orthoses, traditionally or contemporary, and requires future research. Though these practices were indicated for good reliability: Relaxed standing or sitting and partial weight bearing to accurately capture the forefoot and rearfoot width.To model the practicalities of scanning a human foot, a prosthetic lower limb (PLL) was placed on a chair to mimic that of a patient with the foot partially raised. The fat pad expansion and weight could not be replicated at this stage but was considered for further development. The anatomical representation of the foot was collected using a PLL with a white matte finish.

The anatomical representation of the foot was collected using a PLL with a white matte finish. The Artec Leo (Artec 3D, Luxembourg) 3D scanning light reflectance technology created a digital imaging with 3D point accuracy up to 0.1mm and 3D resolution up to 0.2mm at 80 frames per second. The Artec Studio 16 software compiled the point cloud data and created a mesh from the scanned anatomical data. The software allowed for various further post processing stages to create a watertight model.2.2.Parametric insole modellingThe model created in Artec Leo Studio 16 was then exported to Autodesk Fusion 360 (2.0.16265) for further development and construction of the plantar surface. The scanned model informed the development of the insole using a segmented plane through surface and solid modelling for the exportation into nTopology (3.41.2) for the corresponding TPMS to be inputted for the insole design.

The approach offers sophisticated biomimicry and accurate1719d)plantar surface replication which was then assessed to evaluate the potential CAD-informed mechanical and pressure data.2.3.Material and Methods 2.3.1. Gyroid Design and ManufactureTo understand the mechanical behaviour of TPMS structures at insole height for the appropriate lattice design, Gyroid, was chosen to demonstrate our proposed methodology. The upper and lower extremity of the insole were chosen at 10mm height and 5mm height from the approximations gathered from academic research (Ahmed et al., 2020; Bus et al., 2004; Guldemond et al., 2007). Previous research from Claybrook, et al., 2022; Holmes et al., 2022, characterised TPU and TPMS structure with 10mm cell sizes following ASTM D695-15 standard for polymer compression (ASTM International, 2023), this approach was also used in this study. The cells can be examined in a full unit cell and half unit cell; 10mm and 5mm.

The cells can be examined in a full unit cell and half unit cell; 10mm and 5mm. Characterising test samples as 10mm height and 10mm cell (10CH), 10mm height and 5mm cell (10C5H) and 5mm height and 5mm cell (5CH). The strut thickness 1.2mm and 15mm diameter were used as a preliminary pilot for the capabilities of the structure at the different insole heights. The corresponding porosities are shown in Figure 2.a)b)c)Figure 2- Gyroid porosities printed with 1.2mm strut thickness with a) 10mm Height, 10mm Cell, b) 5mm height and 10mm Cell, and c) 5mm Height and 5mm cell. d) a table representing the corresponding porosities with chosen strut thickness, height, and cell size.Structures were fabricated using a Prusa i3 MK3S+ (Prusa Research, Czech Republic) FFF printer, using a shore hardness 90A TPU polymer (Technology Outlet, UK). The Gyroid structures were designed using nTopology to create STL files, prepared for slicing using the ‘generic flex’ profile within Prusa Slicer.

The Gyroid structures were designed using nTopology to create STL files, prepared for slicing using the ‘generic flex’ profile within Prusa Slicer. The structures consisted of 0.2mm layer thickness, print speed 40mm/s, travel speed 60mm/s, nozzle size 0.4mm, 240oC nozzle temperature and 50oC bed temperature.2.3.2. Mechanical TestingCompression testing was conducted in accordance with ASTM D695-15 standard, except the 2:1 ratio guideline, to understand Gyroid behaviour at insole height. Tests were conducted on a 3366 universal test rig mounted with a 10kN load cell (Instron, UK). Samples were tested using an adjusted protocol described previously by Claybrook et al, 2022, where samples were1720maintained at 15mm diameter and two different heights of 10mm and 5mm. The test comprised of 7 repetitions of samples at a 1.2mm strut thickness compressed at a 1mm/min strain rate.3.

The test comprised of 7 repetitions of samples at a 1.2mm strut thickness compressed at a 1mm/min strain rate.3. Results and Discussion 3.1.Scanning capabilitiesThe investigation in this study is to understand the ability of the Artec Leo in 3D capturing the anatomy for diabetic patient’s feet to create appropriate prosthetics from patient specific data. The process of rotating around the patient leg, whilst in the seated position, demonstrated the need for the foot to be raised from the floor. In Figure 3a the distance between the scanner and the plantar surface of the foot created limitations in dexterity with the Artec Leo and a nature of invasive boundaries. The digitised PLL (Figure 3b) highlights the compromise of the point cloud topography at the back of the calf and heel of the foot. The heel is vital in insole manufacturing and requires accurate anatomical representation.

The heel is vital in insole manufacturing and requires accurate anatomical representation. Thus, a minimum distance of 40cm from the ground needs to be specified to increase accuracy and complete 360-degree capture of the foot.Figure 3 - a) Artec Leo scanning Prosthetic Lower Limb, b) Raw scanning image from Artec Leo Scans and imported into Artec Studio 16.The expectation of a patient to hold their leg in a fixed partially raised position is near- impossible, raising practicality concerns and uncontrollable micro movements. Similarly, given the need for fat pad expansion, further apparatus is required to counteract involuntary movements and create a platform for the foot to be partially weighted. Salles & Gyi, 2013 created a raised platform with a scanner contained in a box with a clear cut out for the foot to be captured.

Salles & Gyi, 2013 created a raised platform with a scanner contained in a box with a clear cut out for the foot to be captured. As we want to create partially weighted instead of full weight distribution (Farhan1721et al., 2021), the use of a Perspex platform to withstand the force and create a clear image of fat pad expansion can be replicated for the use of handheld 3D scanning.Another consideration is the use of the PLL is a matte white surface. This is captured excellently with the light surface for the scanner, this may cause complications with different skin pigmentation and requires further evaluation in future work. A possible solution to adjust for this consideration is to note key anatomical landmarks and record the length of the foot alongside the scan, to maintain scale.

A possible solution to adjust for this consideration is to note key anatomical landmarks and record the length of the foot alongside the scan, to maintain scale. Therefore, the following specification is needed for handheld 3D scanning applications of the foot is seen in Figure 4.Figure 4- Guidelines to consider for 3D scanning applications for anatomically representing the foot.3.2.Model FabricationScanning with the Artec Leo was used to inform the digital workflow of a heightened custom- made insole approach. The import of the PLL data into the Artec Studio software is shown inFigure 5- The fill holes option on Artec Studio 16. The use of manual bridging for increased control and accuracy.1722Figure 3b. The software’s ribbon was followed to understand the capabilities of the filling of data, holes and smoothing of the model while maintaining accurate anatomy.The model produced with the superior design quality was achieved by bridging (Figure 5).

This provided more control, following the natural curves of the PLL. With respect to intricate detailing such as toes, the scan managed to capture the grooves and surface texture when maintained in sharp fusion but lost some quality with the use of fast fusion. This is not of concern as the plantar surface is only of interest for insole manufacture, therefore fast fusion was optimal for a smoother surface to work with. The general anatomy was maintained with respect to overall characteristics and silhouette, such as arch height, but further research is needed for the effects on real life patients skin textures.When the initial 3D model was imported, a plane was created below the plantar surface. This was then divided into multiple rectangular subdivisions to create vertices that can adapt to all the different curves of the plantar surface. The result of the pull to the surface of the Artec Leo scan resulted in a need to offset from the plantar surface (Figure 6a).

The result of the pull to the surface of the Artec Leo scan resulted in a need to offset from the plantar surface (Figure 6a). Figure 6b demonstrates the rippling and self-inversion that resulted around the organic form of the foot, to avoid this the edges of the insole required smoothing for a cleaner parametric model (Figure 6c).a)b)c)Figure 6 - a) snapping feature of plane to surface in Autodesk Fusion 360, b) Form of the insole surface model after thickening to insole height (5mm), and c) Insole smoothed to solid body for a simple offset of the prosthetic lower limb.The modelled insole was exported to nTopology for tailored unique loading characteristics. Gyroid was chosen for a theoretical concept of the Smart Insole in Figure 7. This showcases the concept of the ‘smart insole’ split into the four main anatomical regions; toes, metatarasal heads, arch, and heel.

This showcases the concept of the ‘smart insole’ split into the four main anatomical regions; toes, metatarasal heads, arch, and heel. Manufacturing is achieveable from a complete digitised process and customisation, including the corresponding required Gyroid and its mechanical properties.

Manufacturing is achieveable from a complete digitised process and customisation, including the corresponding required Gyroid and its mechanical properties. Patient specific peak pressure information would need to be acquired to inform this manufacturing, creating the opportunity to fully customise the insole from the digitising of the patients specific anatomy and corresponding peak pressure regions on the plantar surface for the appropriate Gyroid and its mechanical properties by functional grading of the insole.1723Figure 7 - Custom made insole from Artec Leo prosthetic lower limb scan data creating a smart insole concept controlled by regional compressive properties of the Gyroid structure a) front profile, b) isometric profile and c) side profile.3.3.Cushioning PropertiesElastomeric structures complex deformable nature inhabits an undefined yield point as a clear break or permanent deformation is absent (Claybrook, et al., 2022).

The geometric configuration of Gyroid allows restoration of the scaffold with a prolonged elastic plateau for energy absorption. The designed porosity controls the characteristics of the Gyroid until complete densification, characterising the structure like solid TPU. Figure 8 shows a closer examination of the singular compression cycle for 10CH, 10C5H, 5CH. Buckling was absent across all porosities and cell sizes, creating a uniform collapse and lower unpredictability of the intrinsic structure. Furthermore, the validation of load axial column theory that buckling is eliminated in the less slender column structures.Figure 8a shows the compressive modulus of the tested samples, with 5CH having the highest flexibility and 10CH the lowest. The intrinsic properties of the cell size and wall thicknesses indicate the importance of their designed volume of space.

The intrinsic properties of the cell size and wall thicknesses indicate the importance of their designed volume of space. Even though the porosity of the 5CH is lower and more material deposited, the influence of the TPMS on the flexibility of the structure is more apparent.The compression testing of the samples (Figure 8b, c, d, e) showed a high repeatability with minimal standard deviation. In comparison to Claybrook et al., 2022 Split P samples, the Split P approximation of pressure at 10% strain was 0.2-0.3MPa which is exhibited in 10CH and 10C5H (Table 1). In Figure 8, 5CH demonstrated a lower threshold that then exceeded both structures at higher strain rates of the samples. This suggests the porosity of the volume of space has a heavy influence on the initial collapse of the Gyroid structure, which is then controlled by the designed cell size. As each sample had the exact same strut thickness but due to their volume of space occupied possessed different porosities (Figure 2).

As each sample had the exact same strut thickness but due to their volume of space occupied possessed different porosities (Figure 2). In Figure 8b it shows there is no clear plastic plateau with 5CH unlike the others, proposing the designed porosity was removed from the structure quicker and densification is unclear in this malleable design, hence the higher MPa values >20% strain for 5CH. In Figure 8c, d, and e 10CH has the least variation in the repeated samples throughout the compression test, and 5CH the most; this1724could be due to the smaller detailed 5CH reaching the resolution of the FFF. Similarly, 5CH requires the highest load and 10CH the lowest with the tested 1.2mm sturt thickness.To understand the energy plateau for each strut thickness, Table 1 shows 10CH, 10C5H, and 5CH up to 60% strain to identify the MPa capable.

From this preliminary strut thickness, it can be assumed that 10mm height insole is preferable to 5mm height as it has a wider range of MPa covering DFUS. In future work we hope to examine Gyroid structures over a wider porosities and strut thickness for a more complete mechanical understanding.a)b)c)d)e)Figure 8 - Data from compressive loading of Gyroid 90A shore hardness TPU for 5CH, 10C5H, and 10CH. a) comparison of the compressive modulus of 1.2mm strut thickness, b) comparison of the average 5CH, 10C5H and 10CH stress strain profiles. The stress strain profiles of c) 5mm cell and 5mm height, d) 10mm cell and 5mm height, and e) 10mm cell and 10mm height.1725Table 1 - Summary of the compressive modulus and compressive stress for all tested samples of 90A TPU Gyroid structures at upper and lower insole heights for 1.2mm strut thickness.4.

ConclusionThe work discussed has put a detailed characterisation to the workflow of applicable scanning technology needed for customised insoles. There is a potential of a complete digitised approach to the medical application of insole fabrication, which could replace the current labour intensive and qualitative techniques utilised clinically. The 3D scanning process has shown a high replication of anatomical landmarks of the feet and the suitability of mimicking the plantar surfaces and applying further 3D printing techniques.The use of TPU and Gyroid structures has shown a possible alternative for insole manufacture with the upper and lower insole heights for characterisation in this study. The preliminary data collected has only used one strut thickness but showcases the possible use of Gyroid at 5mm and 10mm height for an alternative to foam.

The preliminary data collected has only used one strut thickness but showcases the possible use of Gyroid at 5mm and 10mm height for an alternative to foam. There is a wider range of geometrical configurations, TPMS, cell size, and porosity to be explored and attained through compression testing to fully understand this new approach to medical grade insole manufacturing. Similarly, the adjustment of the TPMS’ in conjunction with the plantar anatomy to define the insole is in early stages. The curvature of the patient anatomy may create the need for support and will need to be explored further to limit this.1726ReferencesAhmed, S., Barwick, A., Butterworth, P., & Nancarrow, S. (2020). Footwear and insoledesign features that reduce neuropathic plantar forefoot ulcer risk in people with diabetes: A systematic literature review. Journal of Foot and Ankle Research, 13(1), 1– 13. https://doi.org/10.1186/s13047-020-00400-4Bates, S. R. G., Farrow, I. R., & Trask, R. S. (2016).

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperADDITIVE MANUFACTURING OF SI-SIC CERMETS FOR COMBUSTION DEVICE APPLICATIONSP. P. Radyjowski, D. L. Bourell, D. Kovar and J. L. EllzeyWalker Department of Mechanical Engineering, The University of Texas at Austin, Austin, TX 78712AbstractTraditional manufacturing methods for high-temperature devices are time intensive and limited to simple shapes. Additive manufacturing (AM) reduces lead times and opens the design space to more complex geometries. Indirect laser sintering of siliconized silicon carbide (Si-SiC) cermets was evaluated for creating devices compatible with combustion environments. Heat recirculating combustors especially benefit from geometric flexibility. Si-SiC process improvements are presented for the production of cermet combustors.

Heat recirculating combustors especially benefit from geometric flexibility. Si-SiC process improvements are presented for the production of cermet combustors. The effect of flame on the material was studied by directly exposing samples to hot combustion products at 1000°C and 1260°C for 10 hours. Subsequently, three experimental Si-SiC combustors were manufactured and fired to evaluate the practical aspects of cermet applications. Each device was operated for 70 hours under excess-air methane flames with solid temperatures up to 1405°C. The surface oxidation and phase changes were assessed. Operating temperatures between 1200°C and 1350°C reduce damage to the material and give a promise of long-term, high-temperature operation.IntroductionThe efficiency of thermodynamic cycles is fundamentally tied to the temperature difference between hot and cold reservoirs.

The extension of the operating range envelope provided by high- temperature materials improves the thermodynamic efficiency [1]. Hydrocarbon combustion, the most popular heat source, generates temperatures beyond the capabilities of common materials used to contain the reaction unless dedicated countermeasures are implemented. Several factors are important for the high-temperature selection: ultimate maximum use temperature, strength at temperature, oxidation resistance, creep behavior, and ability to withstand thermal shock. The relative importance of those factors depends on the target application [2]. Historically the common choices were high-temperature metal alloys, such as certain stainless steels and nickel-based superalloys. Their popularity stems from good strength and toughness combined with relative ease of machining. Unsurprisingly, high-temperature metal parts fabrication has dominated the space of additive manufacturing (AM) for high-temperature applications.

Unsurprisingly, high-temperature metal parts fabrication has dominated the space of additive manufacturing (AM) for high-temperature applications. General Electric and Siemens extensively use AM processes in gas turbine combustor manufacturing to reduce the part count of complex assemblies, improve performance, and simplify operations [3, 4]. A 2020 NASEM report on advanced technologies for gas turbines identified AM as a high-priority research area crucial for furthering development in the next decade [5]. However, the metal-based materials are limited to approximately 1200°C, which thermally constrains many designs and impacts device performance. A noteworthy example is an intricate combustor cooling design, fundamentally enabled by AM, that purposely increases the heat loss to keep the material temperature within operational limits [6].

A noteworthy example is an intricate combustor cooling design, fundamentally enabled by AM, that purposely increases the heat loss to keep the material temperature within operational limits [6]. This case represents a deliberate efficiency reduction dictated by materials constraints.157Low-stress devices, such as combustors and heat exchangers, do not require high strength. Therefore the goal for high-temperature material in such application is to withstand the highest possible temperature for the longest time [7]. Technical ceramics are promising materials in such uses due to their elevated operational temperatures. However, the implementation of ceramics materials has been restricted due to their brittle nature and complicated shaping processes. The traditional manufacturing methods, such as mold pressing or slurry deposition, have significant limitations on the final part shape.

The traditional manufacturing methods, such as mold pressing or slurry deposition, have significant limitations on the final part shape. Furthermore, complex ceramic shapes are usually avoided due to the great difficulty of associated machining. Therefore intricate multi-component assemblies are required. A range of AM methods could address this issue as the ceramic part would be created in the final, complex shape. For example, Sobhani et al. [8] used stereolithography to produce tailored microporous ceramic media for low emission combustors. Precise control over the macrostructure improved the performance by position-dependent regulation of the effective thermal conductivity. Furthermore, AM's ability to reliably create intricate geometries with millimeter-scale features opens new possibilities for sophisticated designs that can be produced with greater tolerance and less manual labor.This study examines the applicability of AM materials for combustion device applications.

Combustors were chosen due to the utilization of high-temperature chemistry while being limited by material properties and suitable manufacturing methods. Increased operating temperatures used to be achieved either by incorporating bulk-manufactured ceramics that severely constrain possible geometries or by complex metal designs that employed intricate heat transfer patterns to lower the surface temperatures to acceptable levels. AM manufactured material presented in this work provides higher temperature compatibility without sacrificing shape reproduction capabilities. The combustor that significantly benefits from such improvement is the heat recirculating reactor (HRR), which recycles the heat energy from hot products back into incoming reactants. The required heat transfer conditions necessitate narrow passages, non-trivial geometries, and direct exposure of solid to the flame [9].

The required heat transfer conditions necessitate narrow passages, non-trivial geometries, and direct exposure of solid to the flame [9]. HRR devices are severely constrained either by insufficient temperature limits of metals or the restricted complexity of handcrafted assemblies made from pre- made simple shape high-temperature ceramics. AM also increases dimensional tolerances and reduces the time and manual labor required for handcrafted assemblies. The ability to reliably and repeatedly create intricate geometries with millimeter-scale features opens new possibilities for sophisticated designs and innovative applications. The specific HRR type used in this study is a four-channel counterflow layout, originally created by hand in zirconia and SiC ceramics [10]. It showed extended flammability limits when used as a combustor and could also be used as a fuel reformer or a thermal oxidizer.

It showed extended flammability limits when used as a combustor and could also be used as a fuel reformer or a thermal oxidizer. However, further optimizations and industrial applications were limited due to manufacturing difficulties.MethodsIndirect Laser Sintering of Si-SiCIndirect laser sintering with metal infiltration was first developed at The University of Texas at Austin [11, 12] to produce dimensionally stable yet fully dense ceramic parts. The material of choice was silicon carbide (SiC) infiltrated with silicon (Si) metalloid, which results in a ceramic-metal composite called cermet. The high melting point of Si combined with the excellent high-temperature properties of SiC makes this cermet applicable for combustion applications. A detailed process diagram for siliconized silicon carbide (Si-SiC) manufacturing is illustrated in Figure 1. First, the pure ceramic powder is mixed with a temporary binder compatible with the laser sintering (LS) process (Figure 1a).

First, the pure ceramic powder is mixed with a temporary binder compatible with the laser sintering (LS) process (Figure 1a). Only the phenolic binder is activated during the AM step,158which holds ceramic particles together to produce a green part (Figure 1b). The temporary binder used in this study, partially polymerized phenolic resin (novolac GP-5546, Georgia Pacific Chemicals, Atlanta GA, United States), allows multiple remelting, which is a crucial characteristic for the LS process. Upon completion, the green part is retrieved from the powder bed via mechanical methods. Excavation is a delicate operation as the bonding strength of the binder at this stage is very low. Subsequently, the phenolic binder undergoes a thermally induced polymerization (crosslinking) to prevent further melting at higher temperatures (Figure 1c), hence preserving the part shape. Inconveniently, the polymerization occurs at temperatures above 90°C, beyond the resin's softening point.

Inconveniently, the polymerization occurs at temperatures above 90°C, beyond the resin's softening point. As a result, the part experiences shrinkage in the vertical direction and deformation of unsupported sections due to gravity. The influence of thermal processing in the crosslinking process was examined using a set of thin 100×20×2 mm samples bridged over a 38 mm wide unsupported section, as shown in Figure 6.(a)ceramic particle(b)melted resin(c)polymerized(d)carbonized(e)infill metalbindernew SiCloose powdergreen partcrosslinked partbrown partfinal cermetPowder MixLS ProcessBinder CrosslinkingBinder DecompositionMetal InfiltrationFinal PartAM DeviceOvenVacuum FurnaceFigure 1. Schematic of AM Si-SiC process: indirect laser sintering with metal infiltrationThe crosslinked green state is the most suitable form for modification, such as drilling holes or manual wall carvings.

The part then undergoes a burn-out step, where the binder is carbonized into carbon bridges (Figure 1d). This process takes place in the vacuum furnace at a maximum temperature of 900°C. The result is a porous ceramic preform, generally referred to as a brown part. The traditional approach would proceed with high-temperature direct sintering of ceramic particles. While a part produced using this approach would have intrinsic properties of the neat ceramic, it would also experience significant volumetric shrinkage (>10%), with resulting substantial distortions. Instead of direct sintering, the part is infiltrated with a suitable substance to avoid shrinkage (Figure 1e). For the infiltration step, the brown part is surrounded by pieces of pure silicon and returned to the vacuum furnace, where metalloid melts and wicks into the ceramic structure. The favorable wetting interaction of the Si - SiC pair provides the capillary driving pressure further aided by the vacuum environment.

The favorable wetting interaction of the Si - SiC pair provides the capillary driving pressure further aided by the vacuum environment. Uniquely for the Si - SiC system, when molten silicon contacts the carbonized binder residue, more silicon carbide is produced by a reaction between the Si and the carbon. This reaction further increases the ceramic-to-metal ratio of the final part. There is usually no need for machining, and only minimal assembly is required to interface with the external equipment. Since the infiltrant flows into the voids and there is no physical force at play, only a small shape change occurs relative to the brown part.159Direct Flame ExposureThe direct flame exposure experiment was designed to investigate the longevity of Si-SiC cermet under an excess-air flame. The Si-SiC samples were suspended on a Nichrome wire above a ceramic flat flame burner, further described in Smith et al. [13].

The Si-SiC samples were suspended on a Nichrome wire above a ceramic flat flame burner, further described in Smith et al. [13]. The coupon was positioned such that the flame impinged at its bottom surface, as shown in Figure 2. The apparatus maintained a stable temperature and flame position near the sample for the period of the experiment. A single K-type thermocouple measured the temperature at the middle of sample height (≈ 5 mm above the flame) with flow conditions actively adjusted to maintain the desired conditions.(a)(b)Figure 2. Flame exposure apparatus for AM Si-SiC samples: a) side view b) top viewTwo temperature levels were studied representative of practical, lean combustion applications: a lower 1000°C temperature with higher oxygen content and a 1260°C flame closer to a stoichiometric mixture. Both air and fuel streams were simultaneously stopped at the end of the test, allowing the sample to cool via natural convection.

Both air and fuel streams were simultaneously stopped at the end of the test, allowing the sample to cool via natural convection. The experiment was performed in 5-hour increments for a total of 10 hours. After each exposure, the sample was weighed on a precision laboratory scale (±0.5 mg), and the volume was measured using a water displacement method. Furthermore, optical inspection, scanning electron microscope (SEM) imaging, and energy- dispersive spectroscopy (EDS) analysis were performed after 10 hours of exposure to investigate the oxidation damage.Si-SiC CombustorsThe counterflow HRR combustor design was reproduced in AM Si-SiC using the process described in the Indirect Laser Sintering of Si-SiC section above. The internal geometries of created parts are shown in Figure 3. The previous handcrafted version, used as a reference point for new devices, was limited to simple straight channels, translating into the current AM straight channel pattern (Figure 3a).

The previous handcrafted version, used as a reference point for new devices, was limited to simple straight channels, translating into the current AM straight channel pattern (Figure 3a). The freeform capabilities of AM process enabled modifications to channel shape, which opened new unique research possibilities. A study on the effect of convective heat transfer on combustion characteristics was performed using two modified designs: an AM diverging channel (Figure 3b) and an AM split-channel (Figure 3c). All reactors used a 2 mm separating wall, and the narrowest channel width of 3 mm was achieved in the AM split-wall device. Furthermore, the smallest resolved features were 1 mm 45° chamfers and 1 mm radius fillets on the leading edges of the splitter wall in the AM split-channel reactor. The presented160feature sizes are LS machine-specific, with other teams successfully achieving a microfeature reproduction [14].

The presented160feature sizes are LS machine-specific, with other teams successfully achieving a microfeature reproduction [14]. The finished AM Si-SiC parts were instrumented, linked with a precise flow delivery system, and operated with flames present for extended periods of time. Figure 4 shows the AM diverging reactor under a stable operation with external insulation removed.x = −90 mmx = 0potential flame positionx = 90 mm2 mm(a)half-sinewave(b)flame arrestordiverging sectionsmooth transition3 mm(c)diverging inletsplit-channel sectionFigure 3. The internal structure of AM Si-SiC combustors: a) AM straight channel design, b) AM diverging channel design, and c) AM split-channel designFigure 4.

The internal structure of AM Si-SiC combustors: a) AM straight channel design, b) AM diverging channel design, and c) AM split-channel designFigure 4. AM combustion reactor under operation with upper insulation removed, so the containing flame section is visible.161ResultsIndirect Laser Sintering of Si-SiCThe relative size changes compared with the CAD model were tracked through the manufacturing process and compiled in Figure 5. The CAD-green step involving the LS processing showed a 3% dimensional increase in the layer X-Y plane and 2% growth in the cross-layer Z direction. This change is a stable feature of the laser sintering process usually accounted for by the software prescaling that is known once the procedure is established and characterized. The subsequent 1% decrease in X-Y was close to the measurement uncertainty. However, a substantial 4% reduction in the vertical Z direction registered between green and brown stages was associated with the crosslinking step.

However, a substantial 4% reduction in the vertical Z direction registered between green and brown stages was associated with the crosslinking step. Ultimately, the final volumetric change thorough the process is significantly lower than traditional ceramic densification methods (>10%).X [7.3in] X [185 mm]X [1.6in] Y [40.6 mm]Z [0.96in] Z [24.4 mm]l104%e d o m D A C103%102%. t . r .101%w100%r o t c a f e a c Sl99%98%97%CAD - CADCAD - greenCAD - brownCAD - finalFigure 5. Geometry changes along various stages of the Si-SiC process in planar (X, Y) and vertical stacking (Z) directions.The result of the further deformation test during the crosslinking processing is shown in Figure 6. Two approaches were studied: a fast thermal shock and a slow temperature ramp. The shock test (Figure 6a) minimized the polymerization time by directly exposing the part to high temperatures.

Two approaches were studied: a fast thermal shock and a slow temperature ramp. The shock test (Figure 6a) minimized the polymerization time by directly exposing the part to high temperatures. In this case, the bridge collapsed as the polymerization process was not fast enough to counteract the viscous flow. The new approach (Figure 6b) used in this study uses a slow temperature ramp to allow partial polymerization before substantial softening occurs. A stair-step temperature profile approximated a 1.25°C/hour ramp over 40 hours and was followed by a sharp increase in temperature followed by a hold at 250°C. The resulting deformation was substantially lower, although the required time increased significantly. Prior to this work, the shock crosslinking procedure was commonly used in combination with additional support, either external (sand embedding) or internal (low viscosity epoxy resin infiltration).

Prior to this work, the shock crosslinking procedure was commonly used in combination with additional support, either external (sand embedding) or internal (low viscosity epoxy resin infiltration). The slow ramp approach provided a more straightforward alternative with substantially reduced deformation. However, since the deformation is not entirely prevented, as seen for Z direction in Figure 5, additional support might still be required for dimension-critical parts.Besides the deformation, the Si-SiC process is prone to incomplete infiltration, shown in Figure 7a. This is caused by either a locally insufficient amount of infiltrant or blocking of capillary paths by newly created SiC, as described by Song et al. [15].

This is caused by either a locally insufficient amount of infiltrant or blocking of capillary paths by newly created SiC, as described by Song et al. [15]. Therefore, repetitive remelting with additional silicon was required until the desired density and close porosity were achieved.16238 mm(a)] C ° [ e r u t a r e p m e T n e v O160 140 120 100 80160 140 120 100 800SHOCK 1 hour 250 C153045(b)0153045Time [hours]Figure 6. Unsupported bridge deformation in the crosslinking step: a) fast shock method b) 2-temperature approach c) approximated temperature ramp(a)(b)(c)Figure 7. Examples of process imperfections: a) incomplete infiltration, contrast-enhanced b) Si protrusions in a channel of Si-SiC c) Si pooling in a channel of Si-SiCHowever, the amount of infiltrant had to be carefully controlled, as excessive silicon tends to protrude from the part surface after solidification, Figure 7b.

Si pooling can also occur in severe cases of over-infiltration, Figure 7c. The cause was identified by Stevinson et al. [16], who showed that the positive volume change upon solidification of Si causes excess metal to squeeze out of the structure as it solidifies.Direct Flame ExposureThe results of the direct flame exposure experiment are summarized in Table 1. The mass and volume history shows that the initial 5-hour exposure caused a mass loss with a volume gain in both cases. The positive volume change is associated with the formation of less dense SiO2. The observed mass loss was more pronounced at higher temperatures despite a lower oxygen content in the combustion atmosphere. After another 5 hours, the sample tested at 1260°C experienced a small mass gain accompanied by further volume increase, indicating passive oxidation.

After another 5 hours, the sample tested at 1260°C experienced a small mass gain accompanied by further volume increase, indicating passive oxidation. This type of oxidation is beneficial for the long-term material life as the growing oxide layer forms a diffusive barrier that substantially limits further deterioration. This behavior agrees with Ernstberger's findings [17] that Si-SiC experiences passive oxidation after the initial period of163mass loss, even at low oxygen concentrations. Meanwhile, the sample tested at 1000°C continued to lose mass after 10 hours, combined with a net volume decrease. This behavior indicated undesirable active oxidation and progressive deterioration of the material occurring at 1000°CTemperatureTimeMass change (total) ± 0.02%Volume change (total) ± 2%25°C (reference)0 hr0.00%0%1000°C5 hr0.12%+ 6%1000°C10 hr0.14%1%1260°C5 hr0.17%+ 4%1260°C10 hr0.14%+ 12%Table 1.

Summary of the total mass and volume changes due to flame exposure.The outcomes of optical inspection, SEM imaging, and EDS analysis after 10 hours of exposure are compiled in Figure 8. The front surface of the low-temperature part is cracked with a spalling taking place, which explains the recorded mass loss. Furthermore, the SEM image of the 1000°C sample showed angular oxide structures poorly attached to the surface. In contrast, the high-temperature sample contained a smooth and well-integrated SiO2 oxide cover. The difference in SiO2 morphology can be explained by its glass transition temperature of approximately 1200°C. For the sample tested at 1000°C, the oxide behaved as a crystalline solid leading to the surface spalling and active oxidation. Meanwhile, the sample tested at a higher temperature was covered in the amorphous form of SiO2, which bonded well to the underlying cermet.

Meanwhile, the sample tested at a higher temperature was covered in the amorphous form of SiO2, which bonded well to the underlying cermet. The EDS analysis of both surfaces further confirmed a higher oxygen concentration in the sample tested at 1260°C despite the lack of visible damage and a lower oxygen concentration in this flow environment. Meanwhile, the low-temperature sample showed a lower oxygen content because most of the SiO2 was removed from the part by spalling.2 mmEDS analysis of surfaceCO1000°C 10 hours100%80%36.841.02 mm% c i m o t A60%40%6.616.156.61260°C 10 hours20%42.90%Figure 8. Optical (left) and SEM (center) images of samples after a 10-hour exposure at 1000˚C (top) and 1260˚C (bottom).

Optical (left) and SEM (center) images of samples after a 10-hour exposure at 1000˚C (top) and 1260˚C (bottom). The outcome of the EDS analysis is shown on the right.164Si45.046.48.6Therefore, if a lower temperature operation is desired, an intentional passivation step during the manufacturing process could be added to grow a well-anchored protective layer. Such layer should not experience severe spalling issues at lower temperatures and would therefore extend the part lifetime.Si-SiC CombustorsAll three AM combustors operated in a stable fashion through the entire set of experimental activities required to collect the lean combustion performance data. Detailed combustion results are presented elsewhere [18, 19]. Each reactor experienced at least 70 hours of operation with ten complete cool-down and warm-up cycles and remained fully operational. In contrast, previous handmade devices exhibited microcracks in the ceramic due to excessive thermal stresses.

In contrast, previous handmade devices exhibited microcracks in the ceramic due to excessive thermal stresses. The three AM reactors studied showed no structural issues with high thermal gradients, even as the flame impinged on the cold, room-temperature wall during the startup period. Internal surfaces were inspected visually via a borescope video probe at regular time intervals. The surface degradation of the AM straight channel reactor at the section experiencing the highest temperature is shown in Figure 9. The remaining reactors exhibited similar behavior and therefore are not presented. Within the first 5 hours, protrusions appeared in the hottest part of the channel. After 5 hours, the initial stage of protrusion formation was characterized by smooth, metallic skin. By the 10-hour mark, a significant amount of new growth could be observed. At this point, all channels were mechanically cleaned since the protrusion meaningfully obstructed the flow.

By the 10-hour mark, a significant amount of new growth could be observed. At this point, all channels were mechanically cleaned since the protrusion meaningfully obstructed the flow. During the remainder of the experiments (10 to 70 hour period), only the SiO2 oxide spalling was observed with no further protrusion growth. This degradation mode was consistent with the observations made on samples tested at lower temperatures under direct flame exposure, discussed in the previous section.The generation of solid 0.2 - 4 mm chunks within the channels is an issue for practical applications. A prolonged exposure caused further growth into an irregular shape and an opaque appearance. Notably, this occurred only in the hottest section of the channel, yet at temperatures below the melting point of pure Si metalloid (1414°C).

Notably, this occurred only in the hottest section of the channel, yet at temperatures below the melting point of pure Si metalloid (1414°C). A representative sample was recovered and inspected under the SEM, with results compiled in Figure 10.800 - 1100°C1200 - 1400°Ccleanfull range of conditions0 hrs5 hrs10 hrs10 hrs30 hrs60 hrsFigure 9. Wall damage as observed from within a channel in an AM straight combustor. Mechanical cleaning was performed at 10 hours.165(a)(b)(c)Figure 10. Combustion-induced protrusion: a) SEM overview b) SEM view of a crevice in oxide skin c) optical view on 10×10 mm hatch patternThe outer surface is a smooth oxide skin, akin to the observation made on the sample tested at high temperature under direct flame exposure. The cavity in the upper end of the object was a suitable target for EDS. The analysis showed a significant prevalence (>80 at%) of silicon.

The cavity in the upper end of the object was a suitable target for EDS. The analysis showed a significant prevalence (>80 at%) of silicon. Furthermore, an overexposed image of the shaded area of the cavity, Figure 10b, shows surface features characteristic of metallic substance. Rezaeia et al. [20] reported similar silicon bead formation for Si-SiC lattice samples after a 1-hour hold in a 1400°C air atmosphere. According to their findings, the local concentration of impurities in the Si phase depressed the newly formed alloy's melting point below the surrounding conditions. The silicon shrunk upon melting, which facilitated the oxidation within the structure via opened micropores. Subsequently, a formation of sub-surface, low-density SiO2 led to an increase of the internal stresses that squeezed the molten silicon alloy into the surface. Finally, the low wettability of silicon on SiO2 causes beading and coalescence into larger globules.

Finally, the low wettability of silicon on SiO2 causes beading and coalescence into larger globules. The described process may be detrimental to the long-term lifetime of the material since expelled metal leaves open porosity that further accelerates the oxidation process. The formation and spalling of SiO2 were observed in the experimental reactors at the 30 and 60 hours marks in Figure 9. Nevertheless, the surface degradation was minor compared to wall thickness allowing for successful operation throughout the testing period.ConclusionsThe Si-SiC cermet created via indirect laser sintering was investigated for combustion device use. High-temperature and combustion applications are attractive fields for AM manufacturing. The freeform fabrication of parts with operational limits beyond current materials opens new possibilities in the field of combustion. Direct flame exposure tests at 1000°C and 1260°C confirmed the high-temperature compatibility of the studied cermet.

Direct flame exposure tests at 1000°C and 1260°C confirmed the high-temperature compatibility of the studied cermet. The obtained data shows that long-term use of Si-SiC is possible for operational temperatures above the glass transformation temperature of SiO2 (≈1200°C) due to the formation of a passivation barrier. Subsequently, a practical application using a heat recirculating counterflow combustor was166investigated. Three devices were produced in Si-SiC containing various degrees of complexity with the smallest feature size of 1 mm. All devices were tested for at least 70 hours and remained fully operational afterward. The combustion tests performed on AM Si-SiC reactors showed the material suitable for practical combustion applications at temperatures up to 1400°C. The geometric flexibility provided by AM enabled new, unique combustion research opportunities.

The geometric flexibility provided by AM enabled new, unique combustion research opportunities. The wall damage through the experiments showed previously observed SiO2 flakes formation accompanied by Si metalloid beads formation at higher temperatures. This process may be detrimental to the long-term longevity of the material and reduce the material's upper-temperature limit. However, the overall performance of AM-based Si-SiC is favorable for combustion applications and opens new design possibilities.References[1] G. W. Meetham, "High-temperature materials - a general review," J. Mater. Sci., vol. 26,pp. 853-860, 1991.[2] J. R. Davis, Heat-resistant materials, Materials Park, Ohio: ASM International, 1997. [3] D. Appleyard, "Powering up on powder technology," Met. Powder Rep., vol. 70, no. 6, pp.285-289, 2015.[4] R. Huff and T. T. Wohlers, "Redesigned for Additive Manufacturing: Serial production ofa new fuel swirler for Siemens gas turbine," Metal AM, vol. 5, no.

70, no. 6, pp.285-289, 2015.[4] R. Huff and T. T. Wohlers, "Redesigned for Additive Manufacturing: Serial production ofa new fuel swirler for Siemens gas turbine," Metal AM, vol. 5, no. 3, pp. 169-172, 2019.[5] Natl. Acad. of Sciences, Engineering, and Medicine, "Advanced Technologies for GasTurbines," The National Academies Press, Washington, DC, 2020.[6] A. Adamou and C. Copeland, "Experimental and Computational Analysis of Additive Manufactured Augmented Backside Liner Cooling Surfaces for Use in Micro-Gas Turbines," ASME J. Turbomach., vol. 143, p. 071012, 2021.[7] Y. Bar-Cohen, High temperature materials and mechanisms, 1st ed., Boca Raton, Florida:CRC Press, 2014.[8] S. Sobhani, S. Allan, P. Muhunthan, E. Boigne and M. Ihme, "Additive Manufacturing of Tailored Macroporous Ceramic Structures for High-Temperature Applications," Adv. Eng. Mater., vol. 22, p. 2000158, 2020.[9] J. L. Ellzey, E. L. Belmont and C. H.

DITIVE MANUFACTURING OF SOFT AND COMPOSITE PARTS FROM THERMOPLASTIC ELASTOMERSM. Saari, M. Galla, B. Cox, P. Krueger, A. Cohen, E. Richer *Laboratory for Additive Manufacturing, Robotics, and Automation (LAMRA), Department of Mechanical Engineering, Bobby B. Lyle School of Engineering, Southern Methodist University, Dallas, TX 75275AbstractThermoplastic elastomers (TPEs) are low-durometer materials that can support large strains without breaking, making them attractive materials for producing 3-D printed soft components. However, prefabricated TPE filament, especially those with low hardness, cannot be used in typical filament feed extrusion mechanisms that are popular in material extrusion- based 3-D printers today.Therefore, we have developed a mini-screw extruder, small enough to be incorporated on a typical 3-D printer system, and capable of extruding various TPE formulations directly from commercially available pellets.

This paper presents the design and thermal analysis of the mini- extruder, experimental testing of the 3-D printing process for TPEs with nominal hardness in the range of 5 – 52 Shore A, and compression and tension tests of the properties of printed parts. By combining 3-D printing of soft TPEs with rigid thermoplastics, the new system also opens up new possibilities in additive manufacturing of soft and hard composite parts.IntroductionThermoplastic elastomers (TPEs) are materials that can support large strains without breaking (several hundred percent or more) and have good mechanical strength (typical tensile strength 10 – 21 MPa) [1]. TPEs can be readily found in hardness ranges from 95 to 5 Shore A, making them attractive materials for producing 3-D printed soft, rubber like components.Nevertheless, low hardness prefabricated TPE filaments cannot be used in typical filament feed extrusion mechanisms that are popular in material extrusion-based 3-D printers today.

Consequently, 85-98 Shore A is the typical range of hardness for the commercially available products [2]. Other additive manufacturing methods, such as Stratasys Polyjet, use UV- cured liquid photopolymers and can combine rigid and rubberlike materials to simulate nine Shore A values from Shore 27 to Shore 95 [3, 4]. However, the mechanical properties of these materials are usually lacking compared to TPEs, with typical tensile strength in the range of 1.5 – 10 MPa, and elongation at break between 35 – 130%.The paper presents the design and thermal Finite Element Analysis (FEA) of a novel mini-extruder capable of extruding TPE formulations in a wide hardness range, directly from commercially available pellets. The mini-extruder has been incorporated on a custom additive manufacturing system and standard specimens for tensile and compression tests have beenfabricated.

The mini-extruder has been incorporated on a custom additive manufacturing system and standard specimens for tensile and compression tests have beenfabricated. The experimental results for the force-elongation measurements and stress-strain curves are presented for three elastomers spanning the hardness range from 5 to 52 Shore A. Lastly, several multi-material parts 3D-printed from soft TPEs and rigid thermoplastics are presented.Extruder DesignIndustrial screw extruders, such as the Killion LX-4, can produce TPE filaments with relative ease. However, prefabricated TPE filaments with low Shore hardness tend to buckle under the axial stress when fed into the hot-end extrusion blocks prevalent in today’s Fused Deposition Modeling (FDM) 3D printers.

In addition, because of their large elasticity the TPE filaments are hard to grip by the feeding rollers or gears [5].Therefore, our goal was to develop an extruder capable of working with TPE formulations in a wide range of hardness, and compatible with a typical 3-D printer system from the size, geometry, and control and operation perspective. Specifically, the extruder was intended to function with commercially available 5 - 90 Shore A elastomer pellets, and to be incorporated in the Fiber Encapsulation Additive Manufacturing (FEAM) printer, recently developed by LAMRA [6].We have selected a smooth bore - single screw configuration for its compactness and ease of manufacturing. The steel barrel is mounted at 45° to vertical and the elastomer pellets are fed distally through a custom hopper 3D-printed from ABS (see Fig. 1a).

The steel barrel is mounted at 45° to vertical and the elastomer pellets are fed distally through a custom hopper 3D-printed from ABS (see Fig. 1a). The hopper can swivel about the extruder longitudinal axis for filling and emptying the pellets, and is equipped with a vibrator that facilitates pellet flow into the feed section of the screw. The Ø9.3×170 mm constant pitch screw (length-to-diameter ratio, L/D = 18) feeds the pellets through the transition (melting) section of the extruder barrel, which is maintained at the set temperature by a Ø2025 mm, 45 W collar heater as seen in Fig. 1b. A Ø620 mm, 75 W cartridge heater is used to control the temperature of the hot end block.

1b. A Ø620 mm, 75 W cartridge heater is used to control the temperature of the hot end block. The temperature is sensed using a bolt-on washer thermocouple (Omega WTJ-8-36) mounted laterally on the hot-end block and is controlled by a CN142 universal temperature process controller (Omega, Stamford, CT).The hot end block is equipped with a vertically oriented Ø500 µm brass extrusion die (Quintessintal Universial Building Devices [7]) that is utilized as a print nozzle (see Fig. 1b). The extrusion rate is controlled using a high torque 23HS30-2804S-PG47 geared stepper motor (StepperOnLine Motors & Electronics, Nanjing, China) connected to a spare axis of the FEAM printer motion controller (DMC4183, Galil Motion Control, Rocklin, CA). The melt feed rate is calibrated based on the screw’s volumetric capacity and the print nozzle diameter.

The melt feed rate is calibrated based on the screw’s volumetric capacity and the print nozzle diameter. Extrusion can be stopped quickly by stopping and driving the motor in reverse for a prescribed angle based on the TPE used in extrusion.The mini-extruder is assembled on the base plate using a bolt-on solid aluminum mounting bracket. The bracket cap is equipped with fins and acts, together with the base plate, as a heat sink. In conjunction with the collar heater this design creates the appropriate temperature distribution along the extruder barrel to insure proper melting of the polymer material at the lower extremity, while the upper regions are maintained cool enough to prevent pelletaggregation in the hopper. The base plate mounts on the existing vertical plate of the FEAM printer Z-stage through a XY fine adjustment mechanism.(a)(b)Figure 1.

The base plate mounts on the existing vertical plate of the FEAM printer Z-stage through a XY fine adjustment mechanism.(a)(b)Figure 1. CAD representation of the mini-extruder with mounting and adjustment system for integration in the FEAM system (a); detail of the hot-end assembly (b).Thermal Analysis of the Mini-Extruder DesignFor the mini-extruder to function properly it is necessary to produce and maintain a relatively constant temperature at the printing nozzle, high enough to produce easy flow of the elastomer through the extrusion die and to insure intra and inter layer adhesion of the extrudate to the printed material. A high temperature gradient is necessary along the extruder barrel such that the elastomer material is transitioned from close to ambient temperature at the feed section to a temperature higher than the TPE’s melting temperature in the transition and metering sections of the screw.

Moreover, the collar heater has to be able to produce sufficient heat to completely melt the elastomer during fabrication.To test the feasibility of the design we constructed a FEA model of the extruder using the Simulation module from SolidWorks 2014 package (Dassault Systèmes SOLIDWORKS Corp., Waltham, MA, USA). The 3-D CAD assembly file was imported as a Thermal Study with global component contact set to “Bonded”. To speed up the analysis the gear motor assembly was excluded from the analysis, its contribution to the temperature of the working sections of the extruder being minor. Material properties were assigned from the software library or according to the manufacturers’ specifications. The temperature of the cartridge and collar heaters was set as “Constant Temperature” at the desired working value of 483 K (210 °C) used for the temperature controller.

The temperature of the cartridge and collar heaters was set as “Constant Temperature” at the desired working value of 483 K (210 °C) used for the temperature controller. Convection and radiation boundary conditions were set for all exposed surfaces with a convection coefficient of 5 W/m2 K and emissivity 0.07. The ambient temperature was set to 293 K (20 °C). The model was discretized using a tetrahedral solid meshwith 15,978 elements, 30,373 total nodes and maximum element size of 10 mm, as shown in Fig. 2a. A Dell OptiPlex 7010 workstation with 3rd Gen Intel Quad Core i5-3470, and 16 GB DDR3, 1600 MHz RAM was used to run the FFEPlus iterative solver to obtain the steady state solution.(a)(b)Figure 2. Solid mesh of the relevant components of the mini-extruder model with 15,978 elements and 30,373 total nodes (a); FEA results for the temperature distribution for 483 K working temperature (b).The results shown in Fig.

2b show the temperature distribution on the components under analysis. The temperature of the hot-end block is relatively uniform with only 2 K (2 °C) variation in its entire volume. The temperature at the printing nozzle was calculated as 481 K (208 °C), which was deemed adequate for reliable adhesion of the extrudate to the previously deposited material. The power required to maintain the working temperature in steady state was 35.25 W, well within the capabilities of the heaters. Distal from the collar heater the temperature gradient along the extruder barrel exceeded 3,000 K/m.

Distal from the collar heater the temperature gradient along the extruder barrel exceeded 3,000 K/m. Thus, the maximum estimated temperature at the pellet hopper was 332 K (59 °C), sufficiently low to prevent material softening as well as pellet aggregation.Fabrication of Specimens for Experimental TestingBased on the design presented in the previous sections, a prototype of the TPE mini- extruder was manufactured and integrated in the FEAM additive manufacturing system under development by the Laboratory for Additive Manufacturing, Robotics, and Automation (LAMRA) in the Department of Mechanical Engineering, Southern Methodist University, Dallas TX [6]. The process is an extension of Fused Deposition Modeling (FDM) in which both a fiber (e.g., metal wire, or non-metallic fiber) and the matrix (e.g., a thermoplastic polymer) are deposited simultaneously (see Fig. 3a).

3a). In this system the printing hardware (nozzle, extruder,wire feeder, etc.) remains fixed in the XY plane while a Z positioning stage allows the deposition of successive layers in the vertical direction, generating a three-dimensional part [6, 8].(a)(b)Figure 3. Photograph of the FEAM additive manufacturing system equipped with the TPE mini- extruder (a); Test specimens manufactured using the mini-extruder for tensile (ASTM D638-02a) and compressive (ASTM D575-91) testing of material properties (b).In order to assess the mechanical properties of the elastomers, two types of tests have been selected: tensile testing using the ASTM’s Standard Method for Tensile Testing of Plastics (ASTM D638-02a) [9] and compression testing following ASTM’s Standard Test Methods for Rubber Properties in Compression (ASTM D575-91) [10]. Accordingly, two types of test specimens were modeled in SolidWorks 2015.

Accordingly, two types of test specimens were modeled in SolidWorks 2015. The parts were exported to STL file format, and KISSlicer 1.4.5.10 x64 (http://www.kisslicer.com) was used to slice the geometry and generate GCode tool paths. The layer thickness was set to 0.254 mm and the trace width at 0.45 mm. The infill was set as solid with 3 skin layer loops, with 45° raster and direction alternating each layer. A custom script was used to convert from GCode to GalilCode, the language the GalilMC DMC4183 controller employed in the FEAM system. The motor controller was operated using the linear interpolation mode with the magnitude of the velocity vector held constant.Three materials, spanning a wide hardness range were selected for the test specimens. Kraton G1643 M and D1161 P are clear, linear triblock copolymers based on styrene and isoprene, with polystyrene content of 20% and 15% respectively (Kraton Polymers, Torrance, California).

Kraton G1643 M and D1161 P are clear, linear triblock copolymers based on styrene and isoprene, with polystyrene content of 20% and 15% respectively (Kraton Polymers, Torrance, California). The third selected material was ChronoPrene 5A (AdvanSource Biomaterials, Wilmington, MA), which is a biocompatible TPE based on styrenic olefinic rubber and hydrogenated isoprene. Table 1 presents the relevant mechanical properties of these materials as specified by the manufacturers.Table 1.

Table 1 presents the relevant mechanical properties of these materials as specified by the manufacturers.Table 1. Manufacturer specified mechanical properties of selected TPEs UnitPropertyKraton G1643 M 52 >600Kraton D1161 P 32 1300Hardness Elongation at BreakShore A %ChronoPrene 5A 5 700-900Tensile Strength Specific Gravity Styrene / Rubber ratio Melt Flow RangeMPa g/cc - g/10min10 0.90 20/80 12.50-25.0021 0.92 15/85 -0.34-0.69 - - 2-26It was found that the Kraton G1643 M and D1161 P TPEs extrude well in the mini-extruder at the working temperature of 483 K (210 °C), and extrusion velocities of 18.2 and 16.4 mm/s respectively. Good quality parts from ChronoPrene 5A were obtained at lower extruder temperature, 468 K (195°C), and lower extrusion velocity (15.3 mm/s). The dimensions prescribed by the standards and the actual dimensions for the fabricated tensile and compression specimens are shown in Tables 2 and 3.Table 2.

The dimensions prescribed by the standards and the actual dimensions for the fabricated tensile and compression specimens are shown in Tables 2 and 3.Table 2. Prescribed and actual dimensions for the ASTM D638-02a Type I specimensDimension, mm Standard Tolerance D1161 PT G1643 MS ChronoPrene 5AL 57 0.5 57 57 57D 115 5 115 114 116LO 165 - 164 164 165WO 19 6.4 19 19 19T 3.2 0.4 3.2 3.4 2.6W 13 0.5 12.6 13 12.8Table 3. Prescribed and actual dimensions for the ASTM D575-91 specimensDimension, mm Diameter Thickness Standard Tolerance D1161 PT G1643 MS ChronoPrene 5A28.6 ± 0.1 28.56 28.66 28.812.5 ± 0.5 12.40 12.47 12.55Experimental TestingBefore testing all specimens were conditioned for 48 hours at 23±2 °C and 50±5 % relative humidity, in accordance with ASTM Procedure A of Practice D 618. The tests were conducted at 23±2 °C and 50±5 % relative humidity.

The tests were conducted at 23±2 °C and 50±5 % relative humidity. For tensile testing the specimens were placed in the grips of a custom testing machine, taking care to align the long axis of the specimen and the grips with an imaginary line joining the points of attachment of the grips to the machine. The distance between the ends of the gripping surfaces, D, is indicated in Table 2. The speed of testing was ~125 mm/min and the testing was stopped when the extension reached 101.6 mm. The load and the extension values were recorded at constant displacement intervals. As seen in Fig. 4a none of the specimens showed a significant linear region and no pronounced toe regionwas observed. In accordance to ASTM D 638-02a, the stress was calculated by dividing the load- extension curve by the original average cross-sectional area of the specimens (Fig. 4b).

In accordance to ASTM D 638-02a, the stress was calculated by dividing the load- extension curve by the original average cross-sectional area of the specimens (Fig. 4b). Since no linear region was present, the secant modulus was calculated by dividing the maximum nominal stress by the corresponding strain. The results are shown in Table 4.(a)(b)Figure 4. Experimental force-displacement curves for tension testing of Type I specimens (a); Engineering stress-strain curves for tension testing of TPE specimens (secant moduli shown with dashed lines) (b).For the compression testing the specimens were placed between the platens of the testing machine and conditioned by applying 10% of the total load in five successive cycles. For the final cycle the machine was stopped at constant intervals of deflection for recording purposes (ASTM D575-91). Figure 5a shows the force-deflection curves for the three materials under testing.

Figure 5a shows the force-deflection curves for the three materials under testing. Pronounced toe regions followed by regions of linear behavior can be observed for all specimens.(a)(b)Figure 5. Experimental force-displacement curves for compression testing of cylindrical specimens (zero- strain point compensations shown in dashed lines) (a); Engineering stress-strain curves for compression of TPE specimens with (least square fits of elastic moduli are shown with dashed lines) (b).Following ASTM 638-02a recommendations, the linear region of the curves has been extended through the zero-force axis, and the intersection points have been selected as zero- deflection points, from which all deformations have been measured. Using the original cross-sectional areas, the stress-strain curves shown in Fig. 5b have been obtained, and the elastic moduli were calculated by linear regression from the experimental data in the linear regions.

5b have been obtained, and the elastic moduli were calculated by linear regression from the experimental data in the linear regions. The results for the three TPEs used in this work are presented in Table 4. Due to their internal structure, with elastic segments linking harder domains, TPEs deform much easier in tension than compression, as demonstrated by the much larger compression modulus compared to the tension secant modulus for each tested material.Table 4.

Experimental tension secant moduli and compression moduli for 3D printed TPE test specimens.MaterialKraton G1643 M Kraton D1161 P ChronoPrene 5ATension Secant Modulus MPa 0.529 0.216 0.047Compression Modulus MPa 4.13 1.80 0.567Fabrication of Composite Parts from Thermoplastic ElastomersNovel 3D printing techniques, such as FEAM, can already combine rigid thermoplastic materials with metallic and other type of fibers, enabling fabrication of fiber reinforced composites and even electromagnetic active components, such as speakers, sensors, and solenoids [6, 8, 11]. By integrating TPE mini-extruder developed in this project in the FEAM system we gained the ability to seamlessly combine elastomers with rigid materials opens a new, and largely unexplored, range of applications for additive manufacturing technologies.

Several proof of concept parts have been designed and manufactured using Kraton G1643 M TPE and BendLay, a clear rigid thermoplastic material popular in 3D printing. Figure 5a shows a multi- material annular part in which a 2.54 mm (10 fabrication layers) core of elastomer has been sandwiched between two BendLay outer plates of similar thickness. This part demonstrates the ability to deposit the two types of thermoplastics in rapid succession. The adhesion at the interface between Kraton G1643 M and BendLay was found to be very good, making these materials mutually compatible for composite part printing.(b) Figure 5.

The adhesion at the interface between Kraton G1643 M and BendLay was found to be very good, making these materials mutually compatible for composite part printing.(b) Figure 5. Annular composite part with an thermoplastic elastomer core sandwiched between rigid thermoplastic plates (a); Capacitive force sensor with elastomer intermediate layer and outer shells that incorporate metallic wire fabricated monolithically by FEAM (b); proof of concept vibration absorber with elastomer annular region (c).(a)(c)Using the same material combination a fully functional capacitive force sensor with elastomer intermediate layer, which improves sensitivity, and outer rigid shells that incorporate metallic wire electrodes was fabricated monolithically by the newly expanded FEAM system (see Fig. 5b). Finally, to test the ability to seamlessly switch between the two types of printing materials in the same deposition layer, a proof of concept vibration absorber was designed and fabricated.

5b). Finally, to test the ability to seamlessly switch between the two types of printing materials in the same deposition layer, a proof of concept vibration absorber was designed and fabricated. This part includes a compliant and dampening annular region of TPE bordered internally by a shaft housing and externally by a rigid mounting flange, both made from BendLay, as shown in Fig. 5c. As in the previous cases the inter and intra layer adhesion between the two materials was excellent, resulting in a flawless monolithically fabricated composite part.ConclusionsWe have designed, analyzed, constructed, and tested a mini-extruder capable of extruding various TPE formulations in a wide range of Shore A hardness directly from commercially available pellets.

The compact size of the device and ease of control allows it to be incorporated on a typical 3-D printer system, adding the ability to fabricate elastomer or composite parts containing both rigid and elastic, rubber like, thermoplastics. Three TPEs, spanning the range of 5 – 52 Shore A hardness have been selected to test the mini-extruder, and standard test specimens for tension and compression have been fabricated. The experimental results show that the printed material has good mechanical properties for both tension and compression. During the tension tests we reached strain exceeding 180% without rupture and without remnant plastic deformation of the parts. The elastic moduli of the 3D printed material were comparable to the values published by the manufacturers of the TPEs under consideration, indicating good fusion intra and inter layer fusion.

The elastic moduli of the 3D printed material were comparable to the values published by the manufacturers of the TPEs under consideration, indicating good fusion intra and inter layer fusion. To our knowledge, the parts fabricated from ChronoPrene 5A have the lowest durometer ever 3-D printed.Additionally by integrating the TPE mini-extruder in the newly developed FEAM additive manufacturing system we were able to produce good quality composite parts seamlessly combining the deposition of thermoplastic and rigid elastomers both in successive layers and inside the same fabrication layer.The newly developed additive manufacturing capabilities open a wide range of applications for free form fabrication, such as soft robotics, testable prototypes of injection- molded rubber products, catheters and other minimally invasive medical instrumentation, soft consumer products, etc.AcknowledgementsThis paper is based on work supported by the National Science Foundation under grant no.

Krueger, Edmond Richer, Adam L. Cohen,(2015). “Multi-material additive manufacturing of robot components with integrated sensor arrays.” Proc. SPIE 9494, Next-Generation Robotics II; and Machine Intelligence and Bio- inspired Computation: Theory and Applications IX, 949404 (June 20, 2015); doi: 10.1117/12.2179507.[9] ASTM Standard D638-02a, 2003, "Standard Test Method for Tensile Properties ofPlastics," ASTM International, West Conshohocken, PA, 2003.[10] ASTM Standard D575-91 (Reapproved 2001), 2001, "Standard Test Methods for RubberProperties in Compression," ASTM International, West Conshohocken, PA, 2001.[11] M. Saari, M. Galla, B. Cox, E. Richer, P. Krueger, and A. Cohen, (2015). “Active DeviceFabrication Using Fiber Encapsulation Additive Manufacturing,” to appear in the Proceedings of the International Solid Freeform Fabrication Symposium, Austin, Texas, August 2015..

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceADDITIVE MANUFACTURING PROCESS DESIGNDaniel Clymer*, Jack Beuth*, and Jonathan Cagan*Department of Mechanical Engineering, Carnegie Mellon University, Pittsburgh, PA 15213AbstractA current issue in metal-based additive manufacturing (AM) is achieving consistent, desired process outcomes in manufactured parts. When process outcomes such as strength, density, or precision need to meet certain specifications, these specifications can be met by changes in process variable selection. However, the changes required to achieve a better part performance may not be intuitive, particularly because process variable changes can simultaneously improve some outcomes while decreasing others.

In this work, the tradeoffs between multiple process outcomes are formalized and the design problem is explored throughout the design space of process variables. User input for each process outcome is considered and the best combination of process variables is found to achieve a user’s desired outcome.I. IntroductionDirect metal AM is a rapidly expanding manufacturing method that has foundapplications within the fields of aerospace, medicine, and others requiring highly complex structures. However, as metal-based AM is a relatively new form of manufacturing, a lack of good process knowledge in the field can cause build results to be more variable than traditional manufacturing methods, thereby limiting its capability. To utilize the full potential of AM as an alternative manufacturing method, it is necessary to have sound understanding about how a part will be affected by its process variable selection.

To utilize the full potential of AM as an alternative manufacturing method, it is necessary to have sound understanding about how a part will be affected by its process variable selection. Careful selection of process variables can be a deciding factor in whether a part meets its performance requirements or not, and can greatly improve part quality.Individual outcomes of the AM process have recently been studied, including microstructure [1], surface finish [2], and residual stresses [3]. Various process variables have been considered, including beam properties [4], scan strategy [5], layer thickness, and particle size [2]. AM processes have been mapped to examine the effect of process variables on melt pool characteristics [1,6]. Much of this current research has given users a better understanding of the available design space in AM, and shown how various process selections can improve properties of interest.

Much of this current research has given users a better understanding of the available design space in AM, and shown how various process selections can improve properties of interest. In contrast, this study examines the tradeoffs that occur when multiple outcomes are maximized in the same build. In a real design, there are numerous outcomes integral to the product’s build and the ability to select process variables that will give balance between all outcomes is greatly desired. The goal in this work is to derive a design method to find the ideal tradeoff in process variables in metal-based AM to deliver the optimal overall part properties.Often in direct-metal AM, the two most significant process variables are the beamabsorbed power (P) and beam travel velocity (V). These variables have the most impact on the rate of heat entering the system, and therefore are the most important for the transient heat11976conduction problem.

These variables have the most impact on the rate of heat entering the system, and therefore are the most important for the transient heat11976conduction problem. In this research, the behavior of several process outcomes (density, yield strength, surface finish, precision, and deposition rate) is predicted within the P-V design space. Two design methods are discussed: one method based on user preference for multiple outcomes that finds optimal processing directions for the user’s case, and one method based on user requirement for each outcome that finds regions of P-V space satisfying the user requirements. As is outlined in several case studies, these results can direct the user to a choice of beam power and velocity that are different than the machine nominal settings, resulting in a better part. Applying process mapping techniques to the P-V design space allows the user to achieve a better set of process outcomes for a given design consideration.II.

Applying process mapping techniques to the P-V design space allows the user to achieve a better set of process outcomes for a given design consideration.II. MethodsThe desired result of this study is to find the best process variable combination based on auser’s input for a number of process outcomes. The process outcomes included in the study are: density, yield strength, surface finish, precision, and deposition rate. These have been included because they are outcomes that are often considered important to part designers. The process variables are absorbed power (P) and beam scanning velocity (V). Through a number of basic assumptions about the AM process, each outcome is tied back to these two process variables, and equations are created that show how each outcome is expected to change throughout P-V space. The outcomes are calculated for type 316L stainless steel, although most of the trends discussed would be similar in other materials as well.1.

The outcomes are calculated for type 316L stainless steel, although most of the trends discussed would be similar in other materials as well.1. AssumptionsIn order to reduce the process variables down to two (beam power and beam velocity), aseries of assumptions about other important process variables were made.a.) Hatch Spacing and Layer Thickness – The layer thickness (thickness of the new powder layer) and hatch spacing (the distance between the centers of the deposited beads) are variables that can be changed on the machine. However, both variables will typically scale with the melt pool size. This is because the layer thickness and hatch spacing need to be smaller than the melt pool depth and width, respectively, in order for all material to be fully melted.

This is because the layer thickness and hatch spacing need to be smaller than the melt pool depth and width, respectively, in order for all material to be fully melted. Once a factor of safety is chosen (say, 50% of the melt pool depth), good choices for the layer thickness and hatch spacing can be determined by the melt pool size.b.) Powder Size – The powder size distribution is known to affect the surface quality of theresulting part [2]. One constraint on powder size is that it cannot have bigger particles than the size of the melt pool, or else some material would not be melted. The other factor to consider is the drastic increase in cost as smaller particle size distributions are used. This means that it is typically much more economical to use the biggest particle size distribution possible for a given melt pool size. Thus, particle size can be bounded above and below by melt pool size.2.

Thus, particle size can be bounded above and below by melt pool size.2. Melt Pool CalculationsIn order to calculate the process outcomes, the melt pool sizes must be predicted throughoutP-V space. Finite element simulations, modeling the laser powder process, were run using ABAQUS software, and predicted melt pool quantities throughout P-V space. This model was similar to the model created by Soylemez [14], or more recently, Montgomery [15], but for stainless steel. The model simulated a distributed heat source scanning over a substrate of19772stainless steel 316L. The other surfaces in the model were maintained at a boundary condition temperature corresponding to the preheat temperature of the process, 373 K. A symmetric boundary condition was imposed on the mid-plane to reduce computation time. Melt pool properties were calculated for six P-V combinations, which were: 50W-200mm/s, 50W- 500mm/s, 50W-1400mm/s, 200W-200mm/s, 200W-500mm/s, 200W-1400mm/s.

Melt pool properties were calculated for six P-V combinations, which were: 50W-200mm/s, 50W- 500mm/s, 50W-1400mm/s, 200W-200mm/s, 200W-500mm/s, 200W-1400mm/s. These 6 points can be used to map out P-V space in terms of the melt pool characteristics. This type of mapping has been observed to closely follow the real behavior throughout P-V space of a direct metal AM-produced material [14]. All of the calculations in the research are done for stainless steel type 316L.3. Process Outcome CalculationsThe process outcomes considered in this study were: density, yield strength, surfaceroughness, precision, and deposition rate. Curves for each of these outcomes were tied back to the experimental results found for type 316L stainless steel.a) Porosity and Susceptibility to Flaws – Achieving full density is typically one of the most important requirements for an additively manufactured part.

The likelihood of porosity is related to how much overlap there is between the scanning melt pool and the top of the previous layer. Note that because the powder layer will shrink by as much as 55% when it is melted [10], the effective layer thickness (the thickness of powder the melt pool has to melt through) will be larger than the actual resulting layer thickness.Spierings et al [11] found a correlation between the energy density supplied to the powder layer and the resulting density of the part. The energy density, given in Eq. (1) takes laser power, velocity, hatch spacing, and layer thickness into account..𝑃𝑃 𝑣𝑣×ℎ×𝑡𝑡𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙The approximate relationship to energy density is given in Eq. (2),𝐸𝐸 − 𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷 =(2) where h is the hatch spacing, t is the layer thickness, and a=1/E-Density. With lower energy density, more flaws are likely to develop in the part.

(2),𝐸𝐸 − 𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷 =(2) where h is the hatch spacing, t is the layer thickness, and a=1/E-Density. With lower energy density, more flaws are likely to develop in the part. Note the dependence on hatch spacing and layer thickness: changing these can help fix porosity problems even with low power settings.,23𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷 = 0.92 + 12.2𝑎𝑎 − 514𝑎𝑎+ 4318𝑎𝑎b) Yield Strength – Additively manufactured austenitic stainless steel has been observed to form a fine microstructure, due to the high cooling rates characteristic of the AM process [12–15]. Although a detailed model has not been created to describe yield strength dependence in processing space, it has been shown that this microstructure gets finer with faster characteristic cooling rates, and that the strength increases with these finer microstructures [14,16].

Because a detailed model governing strength throughout processing space has not been found thus far, this research assumed a Hall-Petch relationship, which says that there is an inverse relationship between yield strength and grain size in a material. This type of relationship occurs because moving dislocations, which cause eventual yield in the material, are impeded at the grain boundaries. This makes materials with smaller grains (more grain boundaries) stop dislocation movement better and therefore take more applied stress before yielding. The well-known Hall-Petch equation is given in Eq. (3):,𝜎𝜎 = 𝜎𝜎0 + 𝐾𝐾𝐻𝐻𝑑𝑑 1978−1/2(1)(3)3where σ is the yield strength, d is the grain size, and σo and KH are constants.c) Surface Finish – Surface finish is typically defined in terms of Ra, which is the average of the absolute value of the profile height deviations from the mean height of a surface [17]. Spierings et.

Spierings et. al [2] showed that the as-built Ra of an AM part scales with the average particle size of the powder in metal-based AM. Assuming a linear scaling (we expect particles half the size to have half the Ra value), the expected trend can be predicted throughout P-V space. The equation used for surface finish is given in Eq. (4):,where SF is the surface finish, D is the average particle diameter, and C is a constant 𝑆𝑆𝑆𝑆 = 𝐶𝐶 × 𝐷𝐷 depending on material.d) Precision – The obtainable precision refers to the smallest feature that the machine is able to reliably deposit during the build. This value scales with the average melt pool width, because the process is only capable of creating features that are bigger than one of the melt pools [18]. The equation used to govern precision changes throughout the domain is given in Eq.

The equation used to govern precision changes throughout the domain is given in Eq. (5):,where P is the precision, MPW is the melt pool width, and C is a constant.𝑃𝑃 = 𝐶𝐶 × 𝑀𝑀𝑃𝑃𝑀𝑀e) Deposition Rate – The deposition rate is the rate at which new melted material is being addedto the part. Eq. (6) gives the equation used to calculate deposition rate:,where t is the layer thickness, h is the hatch spacing (distance between the center of two beads), and v is beam scanning velocity. As discussed in the assumptions section, the layer thickness and hatch spacing should also scale with melt pool size, which can relate the deposition rate back to the melt pool size and ultimately to the power and velocity settings.𝐷𝐷𝐷𝐷 = 𝐷𝐷 × ℎ × 𝑣𝑣III. Results1.

Results1. Design Tool #1: Weighted-Sum MethodAs discussed in section I, the weighted-sum method reduces the multiple objectives (density,yield strength, surface roughness, precision, and deposition rate) down to one objective by a weighted summation of each objective. Because of the differing ranges in process outcomes, each outcome was normalized to have a maximum value of one and a minimum value of 0 in the range of P-V space. The weights were determined by user preference for each outcome, with a user input between 0-10, with 10 being the most important. The resulting weighted-sum objective function was mapped out as a contour plot in P-V space.a. Weighted Sum Method: Case StudyIn an example using the weighted-sum design tool, consider user preferences weighting yield strength and deposition rate highly.

Weighted Sum Method: Case StudyIn an example using the weighted-sum design tool, consider user preferences weighting yield strength and deposition rate highly. The weightings are given in Table 2, and the contour map of the weighted-sum objective function is shown in Figure 7.1979(4)(5)(6)4Table 2: Weighted Sum Case Study RequirementsOutcome Density Yield Strength Surface Roughness Precision Build RateUser Weighting 3 7 1 3 5Figure 1: Weighted-Sum contour plot in P-V space for the weighted-sum case study #2Figure 7 shows two regions of higher contour values, both the upper right and lower right regions of P-V space. In this case, the user could make process improvements by making a move in either direction in P-V space.2. Design Tool #2: Bounded Objective MethodThe second design tool explored in this study took user requirement for all but one processoutcome and found a “processing window” in P-V space that satisfied each of these requirements.

Within the processing window, the final outcome was optimized to direct the user to the best P-V combination within the window.a. Bounded Objective Method: Case StudyAs an example, consider a part that will have small features, requiring tight precisiontolerance, as well as a tight regulation on part density. For illustration, requirements are shown in Table 3. After meeting each of these requirements, it is desired to maximize the build rate. Figure 8 shows a contour plot of build rate in P-V space, overlaid with each constraint requirement. The green (shaded) region shows the area that satisfies all build requirements. In this case, the yield strength constraint is active.

The green (shaded) region shows the area that satisfies all build requirements. In this case, the yield strength constraint is active. The optimum within that region is in the upper right, where the build rate is maximized.51980Table 3: Bounded Objective Case Study #1 RequirementsSurface Finish (Ra) Precision Density Yield StrengthLess Than 25 μm Less Than 50 μm Above 99.5% Greater Than 490 MPaMaximize Build RateFigure 2: Case Study #1, showing objective contours of build rate overlaid on design constraint requirementsA designer for additive manufacturing could use this information to alter the process input variables to achieve a part more closely aligned with their required output. In this example, the optimal P-V combination that was found is: P = 186 W, V = 1400 mm/s. This deviates from the machine nominal settings for 316L stainless steel, which are P = 98 W, V = 1083 mm/s.

In this example, the optimal P-V combination that was found is: P = 186 W, V = 1400 mm/s. This deviates from the machine nominal settings for 316L stainless steel, which are P = 98 W, V = 1083 mm/s. They can also use this to have feedback on which constraints, if relaxed somewhat, would change the optimal solution, and which would not affect the optimum. In this example, relaxing the constraint on yield strength would change the location of the optimum.IV. DiscussionThis study explored two methods of assisting an AM user in finding the best regions of P- V operating space: one method using a weighted summation of objectives, and one method using a bounded objective function method. The first can be used to show the user the best direction to move in processing space by finding the direction of the steepest gradient change.

The first can be used to show the user the best direction to move in processing space by finding the direction of the steepest gradient change. The second can be used in situations that the user has more knowledge of specific outcome requirements, to show the user a “processing window” that satisfies all requirements. Typically, the suggested processing variables will show that a deviation from the nominal machine settings provides a build result more closely aligned with the desired outcomes.In the case study involving the weighted-sum objective function, the P-V contour mapshowed the best regions for the user’s preference in process outcomes. One of the most important uses for this kind of contour map is to show the user the optimal direction to move away from the nominal settings in processing space.

One of the most important uses for this kind of contour map is to show the user the optimal direction to move away from the nominal settings in processing space. This is especially useful for AM users that have a limited understanding of how important process outcomes will change in P-V space, and want to have an idea of what parameter changes will benefit their goals before investing in time- consuming and costly experiments on the machine.In the case study involving the bounded objective function, the results suggested a deviation away from the machine nominal settings, which for these examples is P = 98 W, V = 1083 mm/s. This demonstrates the fact that these process variables can be tailored for each specific situation, to create better design outcomes than would have resulted from the nominal settings, depending on user requirement for that situation.

These examples also demonstrated the concept of a “processing window” that the user could operate within to satisfy their process61981requirements. This kind of design tool is useful for more experienced AM users that know exactly what requirements they need to meet in order to have a successful part.In many cases, the machine nominal settings will fall inside this “processing window”. This is because the nominal settings were likely selected to achieve a decent result for a wide range of requirements. However, for specific situations, these nominal settings will not necessarily be the best choice of process variables.

However, for specific situations, these nominal settings will not necessarily be the best choice of process variables. A process that is designed with user requirement in mind, that varies from case to case, can achieve better outcomes for each case than a machine operating nominally.Because additive manufacturing is such a new technology, coupled with the fact thatmost users do not deviate from the nominal process variables, much work still needs to be done to characterize different AM processes and materials. In this work, general expected trends are used as an example to show how a designer might formulate the problem of process design. It is expected that experimental results would yield somewhat different values, compared to the predictions in this study, but the trends are expected to be similar. In the future, it will be of great use to the designer to have more comprehensive experiments characterizing all of the design space.V.

In the future, it will be of great use to the designer to have more comprehensive experiments characterizing all of the design space.V. ConclusionsOne of the most appealing aspects of additive manufacturing is the ability for builds to be customized specifically to a user’s application. Applying process mapping techniques to the P-V design space in additive manufacturing has produced insight into how an AM user might approach the selection of process variables. This study demonstrates an opportunity to design the process of additive manufacturing to achieve a set of outcomes tailored to the user’s needs. A real part will typically have multiple process outcome that need to be simultaneously improved. In AM, these outcomes often depend on the same input variables, making tradeoffs between design outcomes a necessity.

In AM, these outcomes often depend on the same input variables, making tradeoffs between design outcomes a necessity. Based off of user preference or requirement, design variables can be altered to reach the best combination of process outcomes achievable for a user’s application.In the future, as more research is done to improve the field’s understanding of process outcomes’ relationship to process variables in AM, the assumptions made in this paper will be refined to give more precise results. Additional process variables and processes could be incorporated into the algorithm, resulting in more freedom to the process designer to build an optimal part.

Additional process variables and processes could be incorporated into the algorithm, resulting in more freedom to the process designer to build an optimal part. As knowledge of process outcomes becomes more detailed and can be predicted with more accuracy, additive manufacturing processes will be designed for greater user customization.19827References[1] Gockel, J., and Beuth, J., 2013, “Understanding Ti-6Al-4V Microstructure Control inAdditive Manufacturing via Process Maps,” Solid Freeform Fabrication Proceedings, pp. 666–674.[2]Spierings, A. B., Herres, N., Levy, G., and Buchs, C., 2011, “Influence of the particle size distribution on surface quality and mechanical properties in additive manufactured stainless steel parts,” Rapid Prototyp. J., 17(3), pp. 195–202.[3]Peter Mercelis, J. K., 2006, “Residual stresses in selective laser sintering and selective laser melting,” Rapid Prototyp. J., 12(5), pp. 254–265.[4] Murr, L., Martinez, E., Medina, F., Gaytan, S.

DITIVE MANUFACTURING ROUND ROBIN PROTOCOLS: A PILOT STUDYShawn Moylan1, Joshua Land2, Antonio Possolo11 National Institute of Standards and Technologyi Gaithersburg, MD, USA 2 University of Maryland College Park College Park, MD, USAAbstractAs the number of users of additive manufacturing (AM) steadily increases, and considering their demand for material and process specifications, the need for standard protocols for round robin studies is increasing accordingly. Researchers at the National Institute of Standards and Technology (NIST) have conducted and participated in several AM round robin studies with the aim not only to characterize the AM process, and material but also to improve the understanding of AM round robin studies themselves. One simple study, a pilot round robin study investigating geometric performance of NIST-owned consumer-grade 3D printers, provides excellent examples of typical results and lessons learned.

One simple study, a pilot round robin study investigating geometric performance of NIST-owned consumer-grade 3D printers, provides excellent examples of typical results and lessons learned. While individual printers produced relatively consistent results, there was significant variability between the printers. This variability existed despite best efforts to ensure participants followed consistent procedures in building the test parts. Further, the variability made it apparent that collecting pedigree data from each build was required to draw any conclusions about potential causes of the variability.1. IntroductionAs additive manufacturing (AM) technologies mature, industry is looking for standards [1]. Of primary interest are AM material and process specifications to improve efficiency of purchasing AM parts, receiving AM parts from suppliers, or certifying AM parts for critical applications.

Of primary interest are AM material and process specifications to improve efficiency of purchasing AM parts, receiving AM parts from suppliers, or certifying AM parts for critical applications. Thorough material and process specifications will need variability, repeatability, or reproducibility data in addition to machine parameters and expected material properties. Round robin studies provide exactly this information, and therefore will be a valuable tool to be used throughout the industry.A round robin study is an experimental methodology where tests are performed independently multiple times by multiple participants and the results are analyzed statistically to assess their variability. This type of study is often used.to determine the repeatability and reproducibility of a "process". The process is often a measurement method but can also be the fabrication of an artifact using a well-defined procedure (e.g., an additive manufacturing process).

The process is often a measurement method but can also be the fabrication of an artifact using a well-defined procedure (e.g., an additive manufacturing process). Round robin studies may also be used to verify that results from a new process agree with those from an existing, trusted process.i Official contribution of the National Institute of Standards and Technology (NIST); not subject to copyright in the United States. The full descriptions of the procedures used in this paper require the identification of certain commercial products. The inclusion of such information should in no way be construed as indicating that such products are endorsed by NIST or are recommended by NIST or that they are necessarily the best materials, instruments, software or suppliers for the purposes described.As the national metrology institute for the United States, NIST has significant experience in interlaboratory studies, key comparisons, and round robin studies to compare measurement methods.

However, there was no existing guidance on conducting round robins for AM (at NIST or anywhere in the literature), and there was very little guidance on round robin studies focused on any manufacturing process. To gain insight into conducting AM-based round robins, NIST has undertaken several, parallel round robin efforts. NIST provided funding and participated in a study led by the Edison Welding Institute (EWI) for the Additive Manufacturing Consortium focused on material properties of IN625 [2]. NIST also led a round robin study of the tensile properties of an AM Cobalt-Chromium alloy produced by powder bed fusion processes (report still being drafted). NIST was a participant in a study organized through the International Academy for Production Engineering (CIRP) to study Uniform Numbering System for Metals and Alloys (UNS) S15500 (15-5) stainless steel (report still being drafted).

While the results of these studies, the properties of AM materials and their variability, were certainly important and valuable to the AM industry, NIST’s primary goal was to learn more about how to better conduct such studies.Full scale round robin studies are expensive and very time-consuming, but we were seeking practical feedback in a short amount of time. We conceived the pilot study described in the current report to resemble a round robin study as closely as possible, but to be fully self- contained within NIST and to be completed in a ten-week summer research project. To do this, we took advantage of the growing number of NIST-owned, consumer-level three-dimensional (3D) printers being purchased to aid NIST staff in their research endeavors. In this study, a test part was built three times on seven different NIST-owned 3D printers. The sizes of several features of the test part were measured three times by the same person.

In this study, a test part was built three times on seven different NIST-owned 3D printers. The sizes of several features of the test part were measured three times by the same person. Variations in the sizes of the features were analyzed focusing on between-printer, within-printer, and between- measurement effects. The experimental methods and results are detailed here because they provide important insight and lessons that can apply to full-scale AM round robin studies.2. Experimental MethodsA growing number of NIST staff members use consumer-grade polymer extrusion machines to aid in their research. This study examined the variability of a test part produced by several of these machines.2.1 Test PartThe part designed for this study was small and simple. An octagonal base encompassed five features: two protruding features (one cylinder and one square prism), two negative (indented) features (one square and one circle), and one center through hole.

An octagonal base encompassed five features: two protruding features (one cylinder and one square prism), two negative (indented) features (one square and one circle), and one center through hole. Figure 1 shows a solid model of the test part. The octagonal base was 40 mm across. The circular features each had a diameter of 10 mm, and square features each had a side length of 10 mm. The octagonal shape was chosen to prevent warping, which is more pronounced on parts whose shapes have sharp angles. The part was designed to allow easy measurement with low measurement uncertainty. A quick build was desired because NIST staff were volunteering their time and machines to the study and minimal time and material were thought to be more palatable. This test part could be completed in approximately 20 minutes.Figure 1.

This test part could be completed in approximately 20 minutes.Figure 1. Solid model of the test part designed and built as part of NIST 3D printer round robin.2.2 Manufacturing PlanThe manufacturing plan is the detailed procedure each participant would follow when performing builds. The study coordinator stressed the importance of abiding by the plan to minimize the influence of many of the variables that might impact builds.The manufacturing plan was developed by the study coordinator to address each step and variable encountered in producing a part on the 3D printer and was demonstrated to create the test part successfully on one system before the plan was distributed to the study participants. The plan defined the required material—Polylactide Resin (PLA) was selected with the requirement that the material be sourced from the machine manufacturer. If a user did not have the required material, it was supplied to the user by the study coordinator.

If a user did not have the required material, it was supplied to the user by the study coordinator. The manufacturing plan described the physical setup of the printer to be conducted prior to starting each build. This setup addressed the build platform surface preparation (a fresh layer of blue painters tape) and a detailed procedure for leveling the platform. The build platform was not heated for these tests. A maintenance protocol to clean the printer nozzle prior to each build was defined. Part location and orientation were defined as the center of the platform with the flat bottom face of the test part base flat on the build platform. The machine settings input into the system’s control software were defined in the manufacturing plan and are listed in Table 1.

The machine settings input into the system’s control software were defined in the manufacturing plan and are listed in Table 1. Finally, the plan provided instructions on how to remove the part from the build platform, and how to clean the part and machine.In addition to simply following the manufacturing plan, the participants were asked to fill out a template to document the steps taken according to the manufacturing plan. This document acts as a checklist, reminding users of the steps they must follow in setting up the build and instructing them to record the values they input. It also provided space to record comments regarding the state of the machine, unexpected errors that required action, or problems with the manufacturing plan.

It also provided space to record comments regarding the state of the machine, unexpected errors that required action, or problems with the manufacturing plan. The process control document was returned to the study coordinator along with the built test parts.Table 1: Machine settings described in the manufacturing plan.Setting Nozzle Temperature Layer Height Object Infill Feedrate (while extruding) Number of Shells Raft SupportsValue 230 °C 0.2 mm 10 % 80 mm/s 2 Off Off/None2.3 Participants and Number of SamplesThe original study parameters called for several NIST staff each to build three test parts in separate builds on their respective machines. Seven NIST staff members volunteered to participate. All of the printers used in the study were from the same machine manufacturer, but one system used a different (older) model than the others. Since the manufacturing plan could be followed using either model, this was deemed to be within the scope of the study.

Since the manufacturing plan could be followed using either model, this was deemed to be within the scope of the study. Unfortunately, one participant was only able to deliver test parts made in acrylonitrile butadiene styrene (ABS), and could not produce the parts in PLA. Since it was unclear if the material type would influence the results, this participant’s samples were not included in the analysis. Fortunately one participant was able to deliver an additional set of three test parts from a second machine in her possession. One participant was actually able to build seven test parts. Another participant was able to produce one test part, but found that the builds were failing (for an unknown reason) after that. While these circumstances deviated from the original study parameters, the data analysis accurately accounts for them.

While these circumstances deviated from the original study parameters, the data analysis accurately accounts for them. In total, 23 test parts produced by six volunteers from seven systems were included in the study.2.4 MeasurementThe as-built geometry of the test part was chosen as the characteristic to define the variability of builds from various laboratories. Dimensional measurements were made by one person with one set of handheld digital calipers. The standard uncertainty in each measurement was 0.02 mm. Nineteen measurements were made of various features of the test part (see Figure 2). Each measurement on each part was taken three times. The caliper was zeroed before each measurement to reduce instrument variability.3. Results and Data AnalysisTo quantify variation in the test parts, the measurement data was grouped by feature, and then separated by printer into seven sections.

Results and Data AnalysisTo quantify variation in the test parts, the measurement data was grouped by feature, and then separated by printer into seven sections. Using the program R [3] – a programming environment for statistical data analysis and graphics – an analysis of variance was performed by fitting a mixed effects model [4] to each feature. The mixed effects model expressed the length of a feature as a sum of contributions from a fixed effect and from multiple random effects. The fixed effect is the size of the feature. The model decomposed variability into contributions from the random effects: differences between printers, differences between builds within the same31721814161915Figure 2. The sizes of nineteen features were measured on each test part.printer, and differences of replicated measurements of the same build.

The sizes of nineteen features were measured on each test part.printer, and differences of replicated measurements of the same build. Each of these contributions was quantified by a standard deviation – the larger the standard deviation, the more important the corresponding source of variability.The results of the analysis of variance show that between-printer variability was larger than both within-printer and measurement variabilities for seventeen of the nineteen features. The only exceptions were features 8 and 12 where the measurement variability was largest. Figure 3 illustrates these results. For features 8 and 12, the measurement protocol did not specify a consistent measurement position along the height of the cylinder and prism for each replicate measurement. This indicates that features may have large straightness or parallelism deviation in the z-direction.

This indicates that features may have large straightness or parallelism deviation in the z-direction. Those exceptions aside, printer variability was not only larger, but often much larger (from 4 to 10 times larger) than the variability attributable to the other sources.A box and whisker plot further demonstrates the dominance of between-printer variability on the overall variability of the data. The box and whisker plot for feature 4 is shown in Figure 4. This plot is typical of most box and whisker plots from this study summarizing each feature. In this plot, each set of box and whiskers represents the measured values of one feature in builds from one printer. The dark horizontal bar depicts the median of the measurements, the height of the box encompasses the middle 50 % of the data, and the bottom and top whiskers depict the minimum and maximum measurement values respectively.

It is easy to see that the boxes are rather short compared to the overall range of the data, and that there is not much overlap between the boxes. This leads to the conclusion that the variability is primarily a result of between-printer effects.4. DiscussionOne of the primary goals of this study was to understand how to improve AM round robin studies. Therefore it is important to look beyond the study results and include lessons learned in the study procedures.Mixed Effect Value (mm)00.050.10.150.2Feature 1Feature 2Feature 3Feature 4Feature 5Feature 6Feature 7Feature 8Feature 9Feature 10Feature 11Feature 12Feature 13Feature 14Feature 15Feature 16Feature 17Feature 18Feature 19Between-Printer EffectWithin-Printer EffectBetween-Measurement EffectFigure 3. Plot of variability effects as determined by a mixed effects model.0.25Figure 4.

Plot of variability effects as determined by a mixed effects model.0.25Figure 4. Box and whisker plot summarizing variation observed in measurement of the size of feature 4. Note that printer A produced 7 test parts, printer B produced 1 test part, and the others each produced 3 test parts. The horizontal dashed line is the designed length.The template filled out by each participant with information about each build was more important than originally thought and should be expanded. The original intent was to promote consistency and document that each participant accurately followed the manufacturing plan. The template consisted of several check boxes to indicate that the participant performed key steps in the manufacturing plan, along with a table where the participants could enter identifying information (name, machine, software version, etc.), values input into the machine software, and observations during the build.

When the results demonstrated relatively large variability, a better template could have been more helpful in interpreting the results. More information from the participants—the pedigree data associated with the machine and build—could have revealed where key differences existed and allowed for a more specific or detailed analysis of variance. For example, if the template requested the age of each system, the analysis of variance could have included this as one of the covariates and investigated its relative contribution to the overall variability. This template would be more important with more complex systems (e.g., metal- based AM systems) with more input variables.A clear understanding among all participants about the requirements of the study and their respective capabilities would have helped avoid a waste of resources. For example, a participant in this study prepared test parts in ABS because he could not produce them in PLA.

For example, a participant in this study prepared test parts in ABS because he could not produce them in PLA. Had the participant known that this was a strict requirement, he would have known he could not participate in the study and would have avoided using his time and material. Again, the loss here is minor, but could be considerably more costly in larger studies with more expensive systems and materials. This type of confusion may occur more often with more complex systems where users have different levels of access to the software based on the license packages purchased by individual users.Gauging participants’ experience, capabilities, and access to relevant resources at the beginning of the study may be valuable. A few of the participants in this study were very experienced with 3D printing, while a couple were novices.

A few of the participants in this study were very experienced with 3D printing, while a couple were novices. It is assumed that all participantsaccurately followed the procedures in the manufacturing plan, but it is possible that lesser experienced users unknowingly made mistakes. It may be possible in future studies to provide more oversight to less experienced users. This is mutually beneficial because the new users gain experience and insight while the oversight may help eliminate this potential source of variability. On the other hand, more experienced users may have modified their system, have better maintenance practices, or know “tricks of the trade” that others do not, allowing them to produce more accurate or consistent parts. If these factors are not addressed in the manufacturing plan or captured in the template, they may not be adopted by other participants in the study and remain invisible to the analysis.

If these factors are not addressed in the manufacturing plan or captured in the template, they may not be adopted by other participants in the study and remain invisible to the analysis. It may be beneficial to get more experienced users involved in planning the study to help improve the manufacturing plan and template.The manufacturing plan may need to contain fewer specific values and more detailed procedures. It was thought that requiring each participant to use the same values would result in lower variability, but that may not have been the case. A hypothetical example can be seen in the nozzle temperature setting used in this study. Each participant used a nozzle temperature setting of 230 °C, as prescribed in the manufacturing plan. That temperature setting worked well in the one laboratory where the manufacturing plan was developed. However, temperature and humidity control in these laboratories were (understandably) coarse.

That temperature setting worked well in the one laboratory where the manufacturing plan was developed. However, temperature and humidity control in these laboratories were (understandably) coarse. It may be that if each printer used a slightly different nozzle temperature, one more suitable of the individual system and laboratory, more consistent parts would have resulted. One way to accomplish this while maintaining some consistency would be for the manufacturing plan to provide a detailed procedure for each participant to follow to determine the proper temperature for that system (the resulting temperature would be documented in the template). This concept demonstrates the need for machine performance testing and standardized calibration or compensation methods.5. Conclusions and Future WorkThe major takeaway from the results of this study is that between-printer variability plays a larger role than within-printer variability in the uncertainty of part attributes.

This happened despite best efforts to ensure that participants used the same procedures and machine settings. This leads to the conclusion that there are sources of uncertainty still to be characterized that cause differences between nominally identical printers. One can speculate on the potential causes of this variability, e.g., differences in age of machines, maintenance or calibration history, etc., that were not adequately controlled by the manufacturing plans. However, one of the weaknesses of round robin studies is that it is often difficult to pinpoint a specific source of variability, especially when inadequate pedigree data is collected along with the test pieces. Since the round robin was intended to be very simple in design, it was impossible to gauge the effects of possibly influential factors.Because AM is a relatively young technology and part variability tends to be large, it may be more important to examine the causes of variability.

If so, then it should best be done in a suitably designed experiment where potential, likely sources of variability are varied deliberately and systematically. With a careful, well-thought-out design, it is possible to conduct such a study including multiple participants. In fact, results from this study and other AM round robin studies that NIST has been involved in show that between-participant variability is relativelyhigh, indicating that including multiple participants is an important component to the study. Future AM interlaboratory studies at NIST will follow this “collaborative experiment” approach.Additionally, the need for round robin studies to characterize AM variability will certainly remain. Users can benefit from these early experiences and lessons learned.

Users can benefit from these early experiences and lessons learned. Toward this end, a new work item in ASTM Committee F42 on Additive Manufacturing Technologies [5] has been established for a “New Guide for Conducting Round Robin Studies for AM.” A standardized approach to round robin studies will encourage broader participation in future studies, clarify requirements for material and process specifications, and simplify interpretation of results.AcknowledgementsThe authors would like to thank the NIST participants of the study who took the time to help a summer intern with their time and resources. Without their help, this round robin would not have been possible.

Without their help, this round robin would not have been possible. Matthew Staymates, Lourdes Bobbio, Stacy Bruss, Chad Boutin, Mark Esser, John Hettenhouser, Zeeshan Ahmed, Arec Jamgochian, Alden Dima, Sean McGivern, and Ryan Shorts all provided valuable contributions to this study.References[1] Measurement Science Roadmap for Metal-Based Additive Manufacturing, 2013,http://events.energetics.com/NIST- AdditiveMfgWorkshop/pdfs/NISTAdd_Mfg_Report_FINAL.pdf (last accessed August 25, 2015).[2] Kelly, S., 2014, “Development of Mechanical Property Data for Laser Powder Bed Fusion Additive Manufacturing of Nickel Alloy 625,” http://ewi.org/eto/wp- content/uploads/2015/04/70NANB12H264_Final_Tech_Report_EWI_53776GTH_Distri bution_Vol_1.pdf (last accessed August 25, 2015).[3] R Core Team, “R: A Language and Environment for Statistical Computing”, R[4][5]Foundation for Statistical Computing, Vienna, Austria: www.R-project.org. Pinheiro, J. C. and Bates, D.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAdditive Manufacturing Utilizing Stock Ultraviolet Curable SiliconeDaniel A. Porter1, Adam L. Cohen1, Paul S. Krueger1, and David Son21 Dept. of Mechanical Engineering, Southern Methodist University Dallas, TX 75206 2 Dept. of Chemistry, Southern Methodist University Dallas, TX 75206AbstractExtrude and Cure Additive Manufacturing (ECAM) is a method that enables 3D printing (3DP) of common thermoset materials. Ultraviolet (UV)-curable silicone is an example of a thermoset material with a large number of industrial and medical applications. 3D printed silicone prototype parts are obtained using a custom high pressure ram, valve, and UV exposure system.This paper will address issues with printing stock UV curable silicone such as electrostatic repulsion, in-nozzle curing, and extrudate slumping.

One solution that addresses two issues is adding carbon black (CB) to the mixture to reduce electrostatic repulsion while also inhibiting UV cure depth, hence preventing material from curing in the nozzle. Evidence shows that too much carbon black can be detrimental to the structural stiffness of the resulting part.IntroductionAdditive manufacturing (AM) is a lucrative and active research area with significant growth potential due to its myriad of commercial applications [1]. Common AM process materials such as acrylonitrile butadiene styrene (ABS), polylactic acid (PLA), polyamide (nylon), and photosensitive polymers are better understood, with more research attention focusing on the expansion of the viable AM material repertoire.

3D printing silicone is the leading edge of research in academic and industrial environments.Polyurethanes, epoxies, urea formaldehyde, polyimides [2], and silicone rubber are thermoset materials that are difficult to 3D print via traditional additive manufacturing methods. Such techniques like material extrusion (ME) which include fused filament fabrication (FFF) are well researched for thermoplastics but are still being investigated for thermosets. Once crosslinking begins in the curing process for these materials, plastic reflow is impractical if not impossible. Introduced complexities may include unwanted cured material accumulation near the nozzle or part, system jams due to material curing in an improper place, mixing and system loading processes, and cleanup and maintenance after a print.

Some benefits of using a ME setup for printing thermosets is the ability to utilize multiple materials (which is not compatible with stereolithography (SLA) and digital light processing (DLP)) not needing to encase the entire print bed to protect the curable material from initiators such as UV, and minimizing waste.Although most materials that can be used in popular ME or FFF are of a rigid plastic nature, recently, advanced elastomers have been investigated in the field of additive manufacturing. Materials such as thermal plastic elastomers (TPE) [3] and silicones [4] [5] are now of interest. Silicone as a printing material for additive manufacturing is one of the newest additions to the 3D1771printing field due to the numerous applications such as medical implants, component insulation, and prosthetics for their chemical, temperature, and moisture resistances along with mechanical flexibility, and weatherability [6] [7] [2].San Draw Inc.

(San Jose, CA) is able to print silicone in full color, adjustable hardness, and with multi-material capability [8]. San Draw Medical (San Jose, CA) is currently printing silicone seals for fuel cells [9]. Sterne Elastomere (Cavaillon, France) is revealing a silicone 3D printer that prints 0.25 mm layers of UV cured silicone [10]. Keyence, a Japanese sensor and printer manufacturer, is also developing a silicone 3D printing [11]. Stamos + braun Prothesenwerk (Dresden, Germany) is making headway into the biomedical field by 3D printing silicone inserts and bone structures [12]. Wacker Chemie (Munich Germany) announced the world’s first industrial silicone 3D printer, the ACEO Imagine Series K [13], which utilizes drop on demand type printing for UV silicones.

Wacker Chemie (Munich Germany) announced the world’s first industrial silicone 3D printer, the ACEO Imagine Series K [13], which utilizes drop on demand type printing for UV silicones. Fripp Design (Sheffield, U.K.) has the Picsima process which uses a needle to dispense a silicone component into a bath containing the other crosslinking component.As of yet, UV curable silicones are not highly biocompatible, but the vast amount of non- medical silicone applications makes investigating additive manufacturing in this area still highly viable. Some polymer materials that are at least somewhat biocompatible include medical grade polyurethanes [14], polypyrrole [15], poly-ε-caprolactone [16], and silicones [17].

Some polymer materials that are at least somewhat biocompatible include medical grade polyurethanes [14], polypyrrole [15], poly-ε-caprolactone [16], and silicones [17]. Other recent advances transparent glass [18], polyvinylidene fluoride (PVDF) [19], and ultra-soft thermoplastic elastomers (TPE) [20]; fiber encapsulation additive manufacturing [21], printing of drug delivery devices [22], stretchable embedded sensors [23], and capacitive force sensors [24].in additive manufacturinginclude printing opticallyIn this work Extrude and Cure Additive Manufacturing (ECAM) is performed at the Laboratory for Additive Manufacturing, Robotics, and Automation (LAMRA) to cure materials that are extruded and cured layer by layer or in situ. High pressures and friction inside the system create some difficulties for the given ME approach.

High pressures and friction inside the system create some difficulties for the given ME approach. Some of the greatest challenges are electrostatic repulsion, an effect that causes two masses to push apart due to significant charge of the same sign, and nozzle jamming. Carbon black is added to the UV curable silicone mix to reduce the repulsion effect and the amount of nozzle jams while curing in situ. The mechanical properties of 3D printed dog bones with different concentrations of carbon black are quantified and discussed in these proceedings.Experimental MethodsMachine SetupThe custom ECAM machine built at LAMRA labs is shown Figure 1. A high-torque geared stepper motor drives a linear sled which in turn operates a high pressure hydraulic cylinder, Figure 1a). Material is loaded into the cylinder by pulling vacuum and in the case of high viscosity, assisted with a pressurized device.

Material is loaded into the cylinder by pulling vacuum and in the case of high viscosity, assisted with a pressurized device. A manual valve at the end of the cylinder allows the material to be pushed through a Teflon high pressure hose, then high pressure metal tubing, and subsequently into a high pressure stainless steel pneumatically actuated valve shown in Figure 1b). From the pneumatic valve, material is deposited and positioned by the X and Y linear stages. The1772constructed 3D printer with UV shielding is shown in Figure 1c). For silicones with additives, a filter is placed between the hydraulic cylinder and the metal plumbing on the Z stage.Figure 1. 3D printer utilized in this research. a) High pressure ram rendering, b) rendering of 3D printer, and c) constructed printer with valve controller and UV exposure system.A custom UV wand holder is 3D printed in ABS and a render of the setup is shown in Figure 2.

This custom holder ensured repeatability of prints by keeping the angle and distance of UV exposure wands fixed. The UV source is a Dymax Bluewave 200 (40 W/cm2 total intensity) which has four wands optically coupled to the source. Four wands covered the immediate part being printed and ensured a cured layer for moderate part size despite extrusion direction. Brass nozzles with short throat lengths are used to ensure the pressure drop in the system is minimized.Figure 2. Custom UV wand holder and high pressure pnuematic valve setup with shield ring.Initial ObservationsInitial prints revealed problems such as significant slumping of certain materials, nozzle material accumulation (Figure 3b)), in nozzle curing (Figure 3a)), and electrostatic repulsion. Slumping was only a problem if the UV curable silicone was not cured immediately after extrusion. One way to get around the slumping problem was to print low aspect ratio beads.

Slumping was only a problem if the UV curable silicone was not cured immediately after extrusion. One way to get around the slumping problem was to print low aspect ratio beads. However this technique simultaneously increased cured material accumulation on the nozzle and interfered with the print process by rubbing against the part, reducing print quality and reliability. It became apparent that curing in situ was a preferred method when using low viscosity, non-shear-thinning materials. The material, if optically translucent, may act as a light guide and gradually cure material in the nozzle thus clogging the system.1773Figure 3. a) Rendering of a typical silicone nozzle clog in which the material cures in the dispense orifice and b) an image of nozzle material accumulation.Electrostatic repulsion is observed in this system configuration with silicones from Momentive Performance Materials (Waterford, NY).

The effect seems greater in higher viscosity materials like UV Silopren 2030 when compared to UV Electro 225-1 which has viscosities of ~450 Pa-s and ~70 Pa-s respectively at 10 s-1 according to Momentive’s technical data sheet. After a few minutes of material extrusion, the silicone upstream of the nozzle starts to locally accumulate electrostatic charge. Combined with the charging that occurs during UV curing, the resulting effect is a significant repulsion force acting on the extruded silicone. This can be observed by the extrudate deviating from its intended deposition path, degrading part geometry and potentially mechanical quality to suffer. Illustrating this effect is Figure 4d) and f). The silicone used in Figure 4 is UV Silopren 2030 and is shown as a free stream in Figure 4a). Observe the lack of an electrostatic repulsion effect to a black nitrile glove in Figure 4b), an uncured blob of silicone in Figure 4c), and an aged 3D printed silicone part in Figure 4e).

Observe the lack of an electrostatic repulsion effect to a black nitrile glove in Figure 4b), an uncured blob of silicone in Figure 4c), and an aged 3D printed silicone part in Figure 4e). Contrast this to a freshly 3D printed silicone part in Figure 4d) and cured silicone blob in Figure 4f). Diversion of the viscous fluid indicates electrostatic repulsion effects.Figure 4. Repulsion effects for a) free stream of UV Silopren 2030 next to b) a nitrile glove, c) an uncured blob of silicone, d) a recent 3D printed silicone part, e) an aged 3D printed silicone part, and f) a bulk UV cured blob of silicone.Reducing the repulsion effect would require eliminating or neutralizing the surface charge of the part, the tribocharged silicone emerging from the nozzle, or both. Adding electrically conductive fillers can ameliorate this problem.

Adding electrically conductive fillers can ameliorate this problem. Options for electrically conductive fillers include antimony-tin doped oxide [25], aluminum, silver, copper, silicon-carbide [26], ZnO [27], carbon black (CB) powder [28], carbon nanotubes [29], among others. The carbon black option is inexpensive and simple but also increases the opaqueness of the silicone, thus impeding the curing1774depth. Impeding the curing depth can be detrimental to part mechanical properties but could also alleviate some problematic observation like silicone curing in the nozzle.CB in silicone has been explored before with research being done to understand the benefits and effects [30] [28] [31] [32]. Electrical conductivity percolation thresholds for carbon black in a medium vary significantly, but some typical values are 10-25 percent by volume when in a styrene butadiene rubber (SBR) [33].

Electrical conductivity percolation thresholds for carbon black in a medium vary significantly, but some typical values are 10-25 percent by volume when in a styrene butadiene rubber (SBR) [33]. The objective is to not make the silicone as conductive as a bulk material, but rather to reduce the localized tribocharging which appears to cause the electrostatic repulsion. With this in mind, UV Electro 225-1 is mixed with small amounts of Vulcan XC605 carbon black from Cabot (Boston, MA). All samples tested and printed were mixed with a FlackTek SpeedmixerTM DAC 150.1 FVZ-K (Landrum, SC) to ensure thorough dispersion of the CB. The samples were mixed at 2500 RPM for 150 seconds, then the composite was scraped off the inner sides of the container and mixed again.

The samples were mixed at 2500 RPM for 150 seconds, then the composite was scraped off the inner sides of the container and mixed again. Even with the Speedmixer, there are conglomerates of CB that exist which clog the nozzle and require in line system filtering.Quantifying Slumping of SiliconeUnderstanding how much a material slumps is important to 3D printing. It is vital to know just how the height of a bead changes as a function of time. To obtain this information ten traces of UV Electro 225-1 and UV Silopren are printed 100 mm long at 20 mm/s. A dwell time of about 30 seconds is held in-between each trace giving a total dwell time range of about 300 seconds. Immediately after the ten traces are printed a UV flood exposure is done at max intensity so that all of the silicone lines are cured as fast as possible.

Immediately after the ten traces are printed a UV flood exposure is done at max intensity so that all of the silicone lines are cured as fast as possible. Traces of 0.25, 0.20, and 0.10 mm height are printed and have an equivalent flow rate that would allow for a 0.5 mm wide trace assuming a pill shaped cross section.The glass substrate with the printed silicone lines is taken to a microscope, Olympus (Shinjuku, Tokyo, Japan) BX60, where a marking compound is drawn across a section of each trace to improve measurability. Height measurements are performed by focusing on the glass substrate and then on top of the marked trace while a Fowler (Newton, MA) Digital Indicator to measure the displacement. Extrudate widths are measured using a digital microscope camera from Dino Lite (New Taipei City, Taiwan) AM-7023B. These experiments are repeated three times for each type of silicone and trace height and then averaged.

These experiments are repeated three times for each type of silicone and trace height and then averaged. Figure 5 shows the glass plate with three sets of slumping experiments right after UV exposure.Figure 5. Three sets of cured UV Electro 225-1 traces with a light marking compound applied.1775Mechanical Property TestingDog bone structures adhering to ASTM standard D638 T1 are 3D printed in UV Electro 225-1 with different loadings of CB. For comparison, bulk cast molds of the same dog bone are done for the same material. Print specifications for the samples are shown in Table 1. There are five prepared samples for each dog bone lot.Table 1.

Print specifications for the samples are shown in Table 1. There are five prepared samples for each dog bone lot.Table 1. Printing parameters for dog bone samples.ParameterValueNotesFeed rate (mm/s) Layer height (mm) Extrudate width (mm) Infill % Infill pattern type Infill angle (deg) Outer perimeters UV wand distance (mm) UV wand angle (deg)15 0.25 0.50 100X, Y total vector velocityTrace to trace spacingRectilinear Alternates by 90 degrees each layer45 3 ~15 30From wand to plane surface Measured from verticalThe samples are covered in talc powder and marked so that the gauge length marks (50mm) may easily be observed with a HD camera. Each sample is loaded into a tensile testing apparatus and strained at 500 mm/min until an end displacement of 100 mm is achieved. Pictures are captured intermittently from above while force data is recorded. The tensile testing apparatus with a bulk 0.00 and 3D printed 0.15wt% loaded is shown in Figure 6a) and Figure 6b).

Pictures are captured intermittently from above while force data is recorded. The tensile testing apparatus with a bulk 0.00 and 3D printed 0.15wt% loaded is shown in Figure 6a) and Figure 6b). A single sample from the bulk and 0.00, 0.15, 0.50, and 1.00 wt% CB 3D printed dog bones is shown in Figure 6c) through g).Figure 6. Dog bones with 0.00 wt% and 0.15 wt% CB marked and loaded into the tensile tester. Bulk dog bone with c) 0.00 wt% and 3D printed UV Electro 225-1 dog bones with d) 0.00, e) 0.15, f) 0.50, and g) 1.00 wt% CB loading.1776Hardness of all samples is measured using a Rex DD-4 Type M durometer (Rex Durometers, Buffalo Grove, IL) and with LabVIEW (National Instruments, Austin, TX) data logging software. Measurements are taken on three different positions on the samples and then averaged.

Measurements are taken on three different positions on the samples and then averaged. Hardness is a good indicator of how well cured each sample is.Results and DiscussionCarbon Black Effect on RepulsionDifferent amounts of CB are added to Silopren 2030 and bulk cured under a UV wand for around 3 minutes. The larger the wt% of CB, the thinner the layer of cured silicone in the bulk solid due to the CB attenuating the curing depth of the UV. Repulsion effects for the different wt% loadings in UV Silopren 2030 are shown in Figure 7a) through e). Figure 7f) shows that a freshly printed part in UV Electro 225-1 next to a stream of the same material having the same CB loading has negligible repulsion effects.Figure 7. Effects of a) 0.01 wt%, b) 0.16 wt%, c) 0.31 wt%, d) 1.25 wt%, and e) 2.50 wt% bulk cured CB loaded UV Silopren 2030 on the repulsion effect given a 0.00 wt% CB free stream of the same material.

Effects of a) 0.01 wt%, b) 0.16 wt%, c) 0.31 wt%, d) 1.25 wt%, and e) 2.50 wt% bulk cured CB loaded UV Silopren 2030 on the repulsion effect given a 0.00 wt% CB free stream of the same material. A 0.15 wt% CB free stream of UV Electro 225-1 f) flows next to a freshly 3D printed silicon part fabricated with the same silicone.In addition to the reduction of the electrostatic repulsion, all prints with carbon black loading of 0.15 wt% or higher showed a significant decrease in the occurrence of nozzle jams that were caused by silicone curing in the dispense orifice. In fact there was only one jam which occurred when 3D printer was left idle for about 30 minutes while the UV exposure system was active. Nozzle material accumulation was lessened with the same amount of carbon black but eventually still occurs with the given setup.Slumping of SiliconeResults for slumping of UV Electro 225-1 and UV Silopren 2030 are shown in Figure 8. From Figure 8a) and b).

From Figure 8a) and b). The slopes of width versus dwell time are greater for the less viscus silicone UV Electro 225-1. Also the heights in Figure 8c) and d) show the same general trend that UV Electro drops faster than UV Silopren 2030. Images of UV Electro 225-1 after 30 seconds of dwell time, Figure 9 a), show widening by 20 percent. 300 seconds of dwell time for UV Silopren 2030 start to show significant widening, Figure 9 b).1777Figure 8. Slumping result widths for a) UV Electro 225-1 and b) UV Silopren 2030. Slumping result heights for c) UV Electro 225-1 and d) UV Silopren 2030.The results of the slumping test show the importance of needing to cure in situ for materials like UV Electro 225-1. These experiments were performed on glass substrates so the results of slumping may be slightly different for a substrate of tacky silicone of the same type. This would reveal the effect of wettability on slumping.Figure 9.

This would reveal the effect of wettability on slumping.Figure 9. Images of a) UV Electro 225-1 trace after 30 seconds of dwell time b) vs UV Silopren 2030 after 300 seconds.1778Stiffness of Dog Bone SamplesInitial dog bone tensile testing revealed that the bulk, 0.00, and 0.15 wt% CB dog bone samples have almost equivalent stiffness responses. The stiffness results are averaged for all five samples in each category for simplicity. The higher 0.50 and 1.00 wt% CB samples had much lower stiffness values as shown in Figure 10a). Secant modulus results for the stress versus strain curves are shown in Figure 10b). A drop in stiffness for higher loaded CB dog bones was expected because UV curing starts to become significantly attenuated enough that the bottom part of the 3D printed traces are not fully cured by the time the entire layer is finished.Figure 10.

Tensile test results for all as cast and as printed dog bone samples.UV Electro 225-1 was found to be fully cured after baking in an oven at 121°C for 10 minutes. Three of the dog bone samples were put into an oven and baked for the same settings to see if stiffness and hardness improvments could be made. The average stiffness response for the three dog bones in the post-oven cured experiment are shown in Figure 11. It is evident from Figure 11 that post-oven curing the bulk and 3D printed samples will fully cure the silicone. An enhancement in stiffness over simple UV curing is also observed with post-oven processing at the prescribed time and temperature.Figure 11. Post-oven cure tensile test results for cast and printed dog bone samples.1779Quantifications for dog bone tensile testing is shown in Table 2 along with the standard deviation for 100% strain measurements.

Once post-oven cured, the samples increased in Durometer like the results of the stiffness tests. Momentive reports that the durometer for a fully cured UV Electro 225-1 sample is about Shore 25 A which correlates to about ~26-27 Shore M (assuming sample thickness is not small) indicating the bulk cast, 3D printed 0.00, and 0.15 wt% carbon black samples are most likely fully cured via UV exposure.Table 3. Durometer results for all dog bone samples. Percent difference is with respect to initial durometer measurements.Hardness, M initialHardness, M post-oven curedBulk Cast 3D Printed, 0.00 wt% CB 3D Printed, 0.15 wt% CB 3D Printed, 0.50 wt% CB 3D Printed, 1.00 wt% CBAverage26.8 26.1 26.1 24.6 22.6Std. Dev. 1.05 1.30 0.31 1.03 1.03Average26.7 26.0 26.5 26.9 27.4Std. Dev. 0.59 0.20 0.17 0.54 0.27% Diff0.48 -0.29 1.55 9.25 21.26Conclusions and Future WorkUV curable silicone dog bone samples are 3D printed with a high pressure ECAM system.

Dev. 0.59 0.20 0.17 0.54 0.27% Diff0.48 -0.29 1.55 9.25 21.26Conclusions and Future WorkUV curable silicone dog bone samples are 3D printed with a high pressure ECAM system. Slumping of silicones is investigated and is found to be significant for the low viscosity and non- shear-thinning UV Electro 225-1 thus leading to the need of in situ curing. Problems with in situ curing using stock silicone materials are in nozzle curing, nozzle material accumulation, and electrostatic repulsion. Carbon black is selected as an inexpensive additive and was found to1780significantly stop system jams due to silicone curing in the nozzle and reduce the amount of nozzle material accumulation.Material property tests revealed that bulk cast, 0.00 and 0.15 wt% CB 3DP dog bones responded approximately the same to applied force. The 0.50 and 1.00 wt% CB loaded 3DP dog bones showed lower stiffnesses indicating that they were not fully UV cured.

The 0.50 and 1.00 wt% CB loaded 3DP dog bones showed lower stiffnesses indicating that they were not fully UV cured. This could be detrimental to printing silicone parts that have high aspect ratios as the entire part could accumulate a large Z height error due to weight. A post-oven cure on three of the samples from each lot increased all of the dog bone stiffnesses to levels comparable to bulk cured material. Thus a small amount of carbon black in UV silicones can allow for continuous high intensity in situ UV curing to minimize slumping effects while retaining mechanical stiffness and harnesses. One downside is that the color of the silicone is tinted.Work to eliminate nozzle material accumulation currently entails using UV opaque materials with smooth surfaces and nonstick properties. The printing process must be optimized because silicone may still cure around the nozzle in a ring shape which could mechanically lock itself in place.

3D printing with polymers: Challenges among expandingPlastics M, Harper CA. Modern Plastics Handbook. Harper CA, editor. McGraw-Hill;Jin Y, Plott J, Shih AJ. Extrusion-based Additive Manufacturing of the Moisture-Cured1781Scott C at 3Dprint.com. 3D Printed Silicone Seals for Fuel Cells Lead to New Service from [9] San Draw. https://3dprint.com/149248/san-draw-3d-printed-seals/; 2016. Last accessed May 2nd 2017. [10] Scott C at 3Dprint.com. Sterne Elastomere Unveils SiO-Shaping 1601 Silicone 3D Printer at K 2016. https://3dprint.com/153427/sterne-elastomere-silicone-printer/; 2016. Last accessed May 2nd, 2017. [11] Buren A at 3Ders.org. 3D printer manufacturer Keyence introduces silicone based 3D printing material. http://www.3ders.org/articles/20161011-3d-printer-manufacturer-keyence- introduces-silicone-based-3d-printing-material.html; 2016. Last accessed May 2nd, 2017. [12] Prothesenwerk.com. http://www.prothesenwerk.com/en/f156000172.html;. Last accessed May 2nd 2017.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperADDITIVE MANUFACTURING WITH MODULAR SUPPORT STRUCTURESIsmail Enes Yigit, Mohammed Isa, and Ismail LazogluManufacturing and Automation Research Center, Koc University, Mechanical Engineering Department, Sariyer, Istanbul 34450, TurkeyAbstractAdditive manufacturing is praised to have low material waste compared to conventional subtractive manufacturing methods. This is not always the case when the computer aided design (CAD) model consists of large overhangs. In such cases, fabrication of support structures are required to fill the space between the CAD model and the manufacturing bed. In post processing, these support structures must be removed from the model. These supports become waste and reduce the buy-to-fly ratio.

In post processing, these support structures must be removed from the model. These supports become waste and reduce the buy-to-fly ratio. In this paper, we present a pre-fabricated reusable modular support structure system which minimizes the fabrication of conventional support structures. The conventional supports are replaced with modular support blocks wherever possible. The blocks are stacked under the overhang with a robot arm until the overhang of the model is reached. Conventional supports can be fabricated on top when needed with fused filament fabrication. This strategy reduces fabrication of conventional supports. Thus, faster fabrication times are obtained with higher buy- to-fly ratios.Keywords: Support structures, Modular, Additive Manufacturing, Robotic1. IntroductionAdditive manufacturing (AM) is a process where the computer aided design (CAD) model is manufactured additively layer by layer [1].

IntroductionAdditive manufacturing (AM) is a process where the computer aided design (CAD) model is manufactured additively layer by layer [1]. The CAD models designed in various CAD formats can be used by the triangulation into a Stereolithography (STL) file [2]. This triangulated mesh is sliced layer by layer to obtain the toolpaths which will be given to the machine for the construction of the physical model. During the construction of the CAD model, it might be necessary to build support structures to support overhanging parts of the model. These support structures, must be fabricated simultaneously with the model, and hence must be accounted for in the toolpath planning [3].The introduction of the support structures increases the overall manufacturing time of the given CAD model. In addition to time, the material cost of the supports needed to create each part must also be taken into account [4].

In addition to time, the material cost of the supports needed to create each part must also be taken into account [4]. These supports will become waste and reduce the buy-to-fly ratio of the given AM process. Since support structures are not part of the final geometry to be created, they need to be removed; a process that requires a significant amount of extra time and effort. Thus, it is in the industries best interest to reduce the amount of support structures required.Several AM methods have been developed to minimize the use of support structures. Some methods try to determine the best orientation for manufacturing with respect to the primary build axis [5]. Others use multiple build axes [6], [7]. Some try to facilitate the easy removal of the support structures [8]. Some have partitioned models into smaller parts which can be fabricated separately to be assembled later [9]. Others use a secondary material for the support.

Some have partitioned models into smaller parts which can be fabricated separately to be assembled later [9]. Others use a secondary material for the support. The use of2142secondary support materials which are either weaker, soluble in a liquid solution, or melt at a lower temperature than the build material was a significant improvement in simplifying the removal of supports. To do this, the extrusion-based equipment should have a second extruder. The secondary material can be extruded in parallel with the current layer of build material [4].This paper proposes a method for fused filament fabrication (FFF) which reduces the fabrication of the conventional support structures by replacing them with pre-fabricated reusable modular support blocks wherever feasible. With the aid of a robotic arm, the blocks are stacked under the overhanging portions of the CAD model until the desired build height is reached.

With the aid of a robotic arm, the blocks are stacked under the overhanging portions of the CAD model until the desired build height is reached. The bulk of the support will be covered with the blocks and the remaining portion can be filled by fabricating conventional supports on top of the stacked support blocks. This strategy reduces the overall fabrication of conventional supports. Thus, faster fabrication times are obtained with higher buy-to-fly ratios.In the next section, the paper starts by explaining the conventional support generation process for the overhanging areas. Section 3 then uses the newly developed modular support structures method to replace the conventional supports wherever possible. Section 4 provides the robot integration for placement of the modular support structures. Section 5 gives examines results with example models: Stanford bunny and the Atatürk portrait. Finally, Section 6 ends the paper with the conclusion of the work.2.

Section 5 gives examines results with example models: Stanford bunny and the Atatürk portrait. Finally, Section 6 ends the paper with the conclusion of the work.2. Conventional support generationGiven the STL file, the overhanging faces of the model that require support structures must be detected. These are found by checking the angle of each facet of the model which is lower than the minimum self-supporting angle from the horizontal XY plane. Such overhanging faces are visualized in Figure 1.𝑘෠𝑖Ƹ𝑗ƸFigure 1. Overhanging faces that require support colored from green to red. Grey faces do not require support.2143Using the vertices 𝑣1, 𝑣2 and 𝑣3 from the STL model, surface normal 𝑁⃗⃗ for each facet iscalculated by taking the vector cross product of two edges of that triangle𝑁⃗⃗ =𝑣12 × 𝑣23 ‖𝑣12‖‖𝑣23‖The Angle from the XY plane is calculated by using cosine law𝜃 = cos−1 𝑁⃗⃗ ⋅ −𝑘෠ ‖𝑁⃗⃗ ‖‖−𝑘෠‖The angle 𝜃 is calculated for each facet.

The minimum self-supporting angle 𝜃 usually ranges from 30 to 60 degrees depending on the AM process and its parameters. Figure 2(a) shows the areas requiring support on the Stanford bunny. Facets where 𝜃 is lower than the minimum self- supporting angle are isolated from the rest of the model. In this list of facets, checking for the edges which are used only once gives us the boundaries of the area that requires supporting structures.After the areas which require support are found, the supports need to be generated. This is done by first sampling the area. First minimum axis-aligned bounding box (AABB) of the model is found. The edges of the bounding box are used to generate two sets of planes with equidistant spacing. These planes face the direction of 𝑖Ƹ and 𝑗Ƹ unit vectors and are perpendicular to the manufacturing bed. The equidistant spacing is determined by the size of the desired support structures inputted by the user.

The equidistant spacing is determined by the size of the desired support structures inputted by the user. The support points are found by first intersecting the boundaries with 𝑖Ƹ planes. The intersection is solved by triangle-plane intersection. This results in line segments lying on the 𝑖Ƹ planes. Then, intersection of these line segments with the 𝑗Ƹ planes give the support points. These are the points which cannot be manufactured without extruding material under it. This space can either reach the manufacturing bed or lie on the model itself.At each support point a ray is cast in the -𝑘෠ direction using Möller–Trumbore ray casting algorithm [10]. If the ray hits a facet, the support starts on the model. Otherwise the support starts from the manufacturing bed. Rectangular support structures are generated using the selected support width and the length from either the model or the manufacturing bed to the support point.

Rectangular support structures are generated using the selected support width and the length from either the model or the manufacturing bed to the support point. These support structures are plotted in green in Figure 2(b).2144(1)(2)(a)(b)Figure 2. Analyzing the Stanford bunny for support structures: (a) Angle of facets less than 90 degrees (d) Conventional support structures generated for each support pointSupport structures that start from the manufacturing bed are appropriate candidates for replacement with modular support structures. Placing the pre-manufactured modular support blocks reduces the deposition of the conventional support structures and thus improves manufacturing time while reducing the material costs. The next section gives the method to replace the conventional supports with the newly developed modular support structures.3.

The next section gives the method to replace the conventional supports with the newly developed modular support structures.3. Modular support structure generationFor reduction in cost and time, it is desirable to replace the conventional support structure with modular support structures. To start, one needs to determine which support structures can be replaced by the modular support blocks. The process starts by dividing the supports into levels in the 𝑘෠ orientation. The height at which the supports are split is determined by the modular block size. Support structures divided into levels are given in Figure 3. Supports which lie on the model are not considered for replacement with the modular block placement, since a flat surface is required to place the modular blocks.2145Support on ModelLevel 4Level 3Level 2Level 1Figure 3. Conventional support structures split into levels with the given modular block height for the Atatürk model.

Conventional support structures split into levels with the given modular block height for the Atatürk model. The boundaries show the areas which require support.For each support structure level, a two-dimensional support map is generated. The map contains the normalized locations of the support structures within the AABB. Each pixel on the map is in the size of one support structure. The map has the value of true for each < 𝑖Ƹ, 𝑗Ƹ > coordinate pair containing support and false for the rest of the coordinates. A sliding window search is implemented on the map to find the optimum place for the modular blocks. This process is shown in Figure 4. The window has the size of a single modular support block. The window first moves in 𝑖Ƹ direction. When the edge is reached, 𝑗Ƹ is incremented and 𝑖Ƹ is swept again until every coordinate pair is searched. At each point, the number of support points covered by the block are recorded.

When the edge is reached, 𝑗Ƹ is incremented and 𝑖Ƹ is swept again until every coordinate pair is searched. At each point, the number of support points covered by the block are recorded. The first block is placed where the maximum number of support points are covered. Then the process is repeated until all the points are covered.(a)(b)(c)Figure 4. Sliding window search for placement of the modular block: (a) Windows is at the origin (b) Window has moved in the x direction (c) Window has found the maximum location2146At the same time the support map is being swept, two other maps are checked. The first map is for the collision with the model and the second map is for collision with previously placed blocks. The model map is given in Figure 5(b) while the block maps are given in Figure 5(c). With these checks, blocks are only placed at feasible locations. A given point on the map may cover more support points, but it will not be selected if there is collision.

With these checks, blocks are only placed at feasible locations. A given point on the map may cover more support points, but it will not be selected if there is collision. After each placement of the block, the covered support points are removed. At the end of the process, if any support point is left which can not be replaced with modular support structures, it is saved for fabrication of conventional support structures. The remaining support points for the Atatürk model are shown in Figure 5(d).(a)(b)(c)(d)Figure 5. Block placement with consideration of collision with the CAD model: (a) Conventional support structure map (b) CAD model map (c) Modular block placement map (d) Remaining support points map.For successful implementation of the proposed modular support structure concept, the design of the modular blocks is critical. The blocks should be placed with precision and should not be2147disturbed during the AM process.

The blocks should be placed with precision and should not be2147disturbed during the AM process. Thus, an interlocking system which limits the movement of the blocks are desired. This can be achieved by using magnetic aligners or mechanical interlocking structures such as that found in the Lego blocks. The next section uses the modular support structure data to formulate the modular support placement strategy with a robotic manipulator.4. Robotic placement of the modular support structuresThe proposed method uses robotic manipulator to place the modular blocks. The block placement for the given models is shown in Figure 6. and the manufacturing setup is given in Figure 7. The blocks are initially stacked on the left-hand side on the stack platform given in cyan. The robot can have a tool changer where two different tools are employed. The tools and the tool changer have not been shown in this work.

The robot can have a tool changer where two different tools are employed. The tools and the tool changer have not been shown in this work. The first tool is a suction cup tool which picks the blocks from the stack and places them at the desired locations. The second tool is an extruder which manufactures the layers of the model and the conventional support structures. The manufacturing bed is given in green, placed in front of the robotic manipulator.Conventional supportsLevel 3Level 2Level 1Modular Support blocks to be place by robotFigure 6. Modular support structures results: (a) The Stanford bunny (b) The Atatürk portraitThe process starts with the extrusion tool depositing the first layers of the model. When the level of the first modular support block is reached, the extrusion process stops. The robot changes the tool to the suction cup tool. This tool is used to pick and place the first level of blocks.

The robot changes the tool to the suction cup tool. This tool is used to pick and place the first level of blocks. After the blocks are placed, the tool is changed back to the extruder. The extrusion continues starting from this level and the process is repeated until all the layers are fabricated. Depending on the overhanging surfaces, new layers are deposited on some of the placed blocks and others are used as a base for new block placement. Given the initial position of the block on the stack, and the2148desired target position on the manufacturing bed, the joint interpolated trajectories for the 6DOF robotic manipulator are generated by using robot toolbox [11]. These trajectories are plotted in red in Figure 7.Robot trajectoryCAD ModelManufacturing bedBlock stackStack platform6DOF RobotPlaced Support BlocksFigure 7.

These trajectories are plotted in red in Figure 7.Robot trajectoryCAD ModelManufacturing bedBlock stackStack platform6DOF RobotPlaced Support BlocksFigure 7. Robotic additive manufacturing setup with modular support block placementWith the introduction of the blocks, the material waste is reduced, and the fabrication time isimproved. The next section gives the results for the proposed method.5. Results and discussionsThe algorithms were implemented on MATLAB© environment installed on a computer with Intel© I7 CPU, 8 GB of RAM. Using Modular support block dimensions of 40 × 40 × 40 [𝑚𝑚] with support spacing of 10 × 10 [𝑚𝑚]. The percent of volumetric improvement in the support structures are calculated and given in the table below.Table 1.

The percent of volumetric improvement in the support structures are calculated and given in the table below.Table 1. Comparison of the proposed method with the conventional methodModel Atatürk portrait Stanford bunnyFacets Vertices Dimensions Conventional Proposed 8.4[cm3] 1527026.3[cm3]7637181x154x226[mm]33398 16701 229x177x226[mm]46.5[cm3]8.0[cm3]Improvement 68%83%With two different examples, it is shown that the proposed method works robustly and canreduce the support structures with significant improvement.6. ConclusionCAD models that consist of large overhangs result in fabrication of large support structures resulting in wasted material and high buy to fly ratios. In this paper, a pre-fabricated reusable support structure system which minimizes the fabrication of conventional support structures is2149introduced.

In this paper, a pre-fabricated reusable support structure system which minimizes the fabrication of conventional support structures is2149introduced. When an overhang is to be fabricated, modular support blocks are stacked under the overhang with a robot arm until the overhang of the model is reached. After the support structure strategy is formulated, The Robot program is generated and integrated into the existing AM G- code. To verify the improvement of the proposed strategy, two models are simulated, and the results are compared to the conventional methods. Minimizing the fabrication of support structures, the material can be reduced by 75% on average, which also results in a reduction in fabrication time.7. AcknowledgementsThis research was supported by the Koç University Manufacturing and Research Center.8. References[1] W. Gao et al., “The status, challenges, and future of additive manufacturing in engineering,”[2][3][4] [5]Comput. Des., vol. 69, pp.

ditive manufacturing with RepRap methodology: current situation and future prospectsL. Romero1, A. Guerrero1, M. M. Espinosa1, M. Jiménez 2, I.A. Domínguez1, M. Domínguez11 Design Engineering Area - Universidad Nacional de Educación a Distancia (UNED), Madrid, Spain2 Department of Mechanical Engineering, Technical School of Engineering - ICAI, Comillas University, Madrid, SpainAbstractIn February 2004, Adrian Bowyer, from the University of Bath, began an open initiative called RepRap, with the purpose of creating an open source rapid prototyping machine which, moreover, could replicate itself. This article analizes the current status of the RepRap initiative, commenting the basic components of RepRap machines, the differences between the different 3D printers developed by the RepRap community so far, and the technical possibilities that opens this technology from the engineering point of view.

In addition we propose some improvements that could be perfectly feasible in the short term. For this purpose, the assembly of a RepRap Mendel Prusa was performed, but with some modifications.1. Introduction to RepRap community.A 3D printing machine (and any computing device in general) is composed of hardware (electric, electronic, electromechanic and mechanic physical components) and control software (the logical equipment as the operating system or software applications), usually both developed by companies that control them.However, there are several companies in the world that are dedicated to the development of hardware in a peculiar way: they publish the sources of their inventions and sharing them with the whole world, with the aim that users from all the planet can use them and propose improvements.

This is the model of open source hardware: hardware whose design, components, tools and documentation are open and they are publicly available to anyone.The origin of these companies is varied: some were linked to projects without profit; others employ directly the open source model; in other cases focus on the research and some educational projects (such as the case of Arduino or RaspberryPi). These initiatives, by the fact of being open, have made possible the emergence of other companies that have been able to develop commercially an original product or offer services and products around that initial product.As in other cases related to free culture, this economy of open hardware is based on the quality of the product and the company innovation capacity and not in the sale of single copies, although it may be under a trademark.

This model seems to work: companies like Sparkfun, Adfruit, Ultimaker, Cooking Hacks or RepRap BCN produce open hardware.RepRap is a rapid prototyping machine, modeled by melting and solidification with a nozzle, self-replicating, and with open source. RepRap is being developed by an open community on RepRap.org, which was founded in 2004 by Adrian Bowyer at the University of Bath. As the term Fused Deposition Modeling is registered by Stratasys, the RepRap community has coined the term Fused Filament Fabrication (FFF) that can be used by anyone without restriction [1]. RepRap is self-replicating because it is capable of generating a significant fraction of its own components.RepRap is registered by the University of Bath, and is open source because it is registered under a 2 GPL versión [2]. Under these conditions, any person can distribute and amend a RepRap machine, but must maintain the modifications under this license.The changes must continue to be publics.

Under these conditions, any person can distribute and amend a RepRap machine, but must maintain the modifications under this license.The changes must continue to be publics. As the machine is free and open source, anyone can build, without the payment of any fee, an unlimited number of copies, for him or any other, using RepRap machines to build the plastic parts of the copies.The founders of RepRap related the prototyping machine with a biological metaphor. They compared it with a mutualism relationship, also known as symbiosis. In this relationship, the machine benefits from the human through its support for playback: the human has to complete the assembly of the new machine with the parts that are not self- replicating and assemble the machine. And the benefit of the human is the use that makes from the machine in order to create the prototypes that he wants.

And the benefit of the human is the use that makes from the machine in order to create the prototypes that he wants. Due to the biological inspiration of their creation, RepRap machines have been baptized with surnames of well-known biologists (Charles Darwin, Gregor Mendel, Thomas Henry Huxley and Alfred Russel Wallace).2. RepRap models and features.3D printers developed by RepRap community have the following characteristics:They are machines of the type FFF (Fused Filament Fabrication), and they use thermoplastic compounds (usually ABS and PLA).Four models have mainly been developed to date (Darwin, Mendel, Huxley and Wallace) with different variants (Prusa Mendel, MendelMax, RepRapPro Huxley, RepRapPro Mendel, etc).The accuracy of the model is given by the diameter of the nozzle, which is usually between 0,4 - 0,5 millimeters.Darwin was the first model developed by RepRap in 2007 (Figure 1).

This first version is currently obsolete.Figure 1: RepRap Darwin.Mendel is the second model developed by the RepRap community, and it is the most popular to date, at least in terms of number of variants. The support structure is a triangular pyramid, formed by M8 threaded rods and printed pieces (Figure 2). The most widespread variant is the RepRap Mendel Prusa, created by Josef Prusa.Figure 2: RepRap Mendel.There is a Mendel variant with the pyramidal structure formed by aluminum profiles (MendelMax) which was created by Maxbots in December 2011. These profiles provide easer assembly and greater rigidity, with a small increase in price.A variant with three extruders has also been developed: the RepRap Tricolour Mendel which was created by RepRapPro Ltd in 2013. The extruders use Bowden tubes which reduce the total weight to support.Huxley is the third model and it was created by RepRapPro (Figure 3).

The extruders use Bowden tubes which reduce the total weight to support.Huxley is the third model and it was created by RepRapPro (Figure 3). It is the small and transportable variant of the Mendel Prusa, using M6 rods and NEMA 14 motors, like the RepRap Wallace.Figure 3: RepRapPro Huxley.Wallace is the smallest model of the RepRap family. It is designed with M6 threaded rods and NEMA 14 motors. It is the most simple and easy to assemble variant (Figure 4). It is an evolution of the Printerbot, the simplest and cheapest RepRap printer.Figure 4: RepRapPro Wallace.3. Basic components of RepRap machines.The basic components of RepRap machines can be classified according to the type of components or the function that they execute.a) Components by type.There are four groups in this category:Plastics components: They are replicable components. They can be manufactured by a RepRap.

They can be manufactured by a RepRap. In Darwin and Mendel machines, the plastic components represent 48% of the total of the elements [3]. The material that forms them is normally PLA (polylactic acid) or ABS (acrylonitrile butadiene styrene). They include structural components, extruder components, carriage print components, etc.Figure 5 includes a picture of a complete set of printed parts. The printer has been designed in order to use metal bearings which are easy to find: circular metal bearings are used by skates’ manufacturers and they are sold as spares in sport equipment storesFigure 5: Complete set of printed partsNon-replicable structural components ("vitamins"): includes threaded and smooth bars, washers, nuts, circular and linear bearings, special parts for the extruder, belts...

They are usually made of metal, steel or aluminum, except the belts that are usually plastics or flexible rubber.Electronic components (“more vitamins"): they are the necessary electronic components for the printer operation, like microcomputer and electronic board, stepper motors, temperature and end stop sensors, resistances and cables.The movement of the machine axes is achieved by stepping motors. The axis of these motors takes a turn with a fixed angle whenever it receives an electrical impulse.

The axis of these motors takes a turn with a fixed angle whenever it receives an electrical impulse. The motors that are often used are the NEMA 17, which is a standard of the National Electric Manufacturers Association (NEMA).Software: The RepRap microcomputer (firmware) and the external computer (for sectioning parts and calculating the GCODE), which are necessary for the equipment operation.b) Components by function.From a functional point of view, the components of RepRap printers can be divided as follows:Support structure: is the structure that supports the whole body of the RepRap.Y, X, Z axes. Y axis movement is forwards/backwards movement; X axis movement is side to side, left to right; and Z axis movement is up and down along the vertical plane.

Y axis movement is forwards/backwards movement; X axis movement is side to side, left to right; and Z axis movement is up and down along the vertical plane. Linear movement is generally accomplished using one of 2 different methods: Belt/pulley driven motion and threaded rod, or lead screw motion.The extruder: is responsible for feeding filament through a nozzle and melting it as it's deposited into the bed where the part is made [4]. The extruder (Figure 6) consists of two parts: The cold end and the hot end.Figure 6: Extruder and extruder parts made by a FFF machine.The print bed: is what the RepRap extrudes plastic into, where the plastic parts are built up. The print bed may be stationary, like the original reprap Darwin, or it may move along one of the x/y/z axes. The bed usually consists of two plates: the upper plate and the lower plate.The heated bed.

The bed usually consists of two plates: the upper plate and the lower plate.The heated bed. While a heated bed is considered an optional component of a RepRap, it often becomes necessary for operating a RepRap long-term because, without a heated bed, parts have a tendency to cool down too quickly.The fusor: melts the fiber and puts it through the nozzle.The end stop: is a very small and simple circuit board with a switch that tells the reprap when it has moved too far in one direction. Thus, there's normally 6 of these: 2 for each axis. A single end stop connects via wires to either: The controller or a stepper driver board.The controller: is the brain of the reprap. It receives orders from outside, processes them and sends them to the printer.Almost all reprap controllers are based on the work of the Arduino microcontroller.

It receives orders from outside, processes them and sends them to the printer.Almost all reprap controllers are based on the work of the Arduino microcontroller. The most used controller are the Arduino Mega (with RAMPS expansion plate), the Sanguinololu (a Arduino variant which includes Pololu controllers and don’t need expansion plate) and the Generation 6 (a design made for the RepRap machines on one plate). All of them have 8-bit microcontroller of the ATMega family.Figure 7 shows a picture of an Arduino Mega [5], with a expansion plate RAMPS v1.4 (Pololu expansion plate of RepRap Arduino Mega).Figure 7: Arduino Mega and plate RAMPS v1.4.The stepper motor: is a type of electric motor that can be accurately controlled by the controller. Most repraps use 4 to 5 stepper motors. 3 to 4 motors control the x/y/z axis movement and 1 motor is used per extruder.The stepper driver: is a chip that acts as a kind of middle-man between a stepper motor and the controller.

3 to 4 motors control the x/y/z axis movement and 1 motor is used per extruder.The stepper driver: is a chip that acts as a kind of middle-man between a stepper motor and the controller. It simplifies the signals that need to be sent to the stepper motor in order to get it to move.The software. There’s specific software used for the machine operation, which can be divided into the following parts:Firmware: is the software installed on the microcontroller, which translates the GCODE received in motor movements, and it controls sensors and heaters. There are several alternatives for the Arduino, the most used being Sprinter [6], Marlin [7] and Teacup [8].GCODE generator: is the program that divides the piece printed in layers and generates the needed GCODE to build them. Installed on the external computer, it transmits the GCODE to the printer via USB or an SD card.

Installed on the external computer, it transmits the GCODE to the printer via USB or an SD card. The most used GCODE generators are Slic3r [9] and Skeinforge [10].Printer control: is the program for basic control of the printer. It allows you to pause and resume printing, set temperatures, control the motors, send GCODE commands and send gcode files with parts to print. The most widely used is the Pronterface [11].The filament. Generally, people use one of two types of filament: ABS or PLA. ABS stinks and warps but is pretty strong, like legos, and PLA smells like waffles and is biodegradable. ABS fumes are also not good for your health. ABS will bend before it breaks and PLA is tougher but breakable as well. So, if you need a gear for instance, use PLA because it will preserve its shape longer.4.

ABS will bend before it breaks and PLA is tougher but breakable as well. So, if you need a gear for instance, use PLA because it will preserve its shape longer.4. User and exchange communities.The social draught of this technology has been enormous and, as we have already said, its development is growing exponentially, allowing that design students and small architecture studies from all the world can test and perfect their creations before building them to real scale.Accessibility to the technology has linked a number of user communities which exchange knowledge and experiences with the purpose of perfecting the printing system.These user communities have developed various exchange platforms to download and print pre-existing 3D models.

Thingiverse and Rascomras are two examples of platforms for loading and downloading of models.Clone Wars is a group within the RepRap community, which seeks to document in Spanish everything in order for anyone to build their own 3D printer [12]. The Clone Wars project seeks disclosure of the RepRap technology, while it brings out new designs and new avenues of research, but not so much in the sense of being self- reproducing. This community was born in 2011 at the Universidad Carlos III of Madrid with Juan González, robotics professor who had previously attended a workshop taught by Adrian Bowyer. At the beginning of 2012 there were only 10 machines or clones; nowadays there are more than 223 clones which more or less actively participate in the community.During the last 3 years, a series of workshops, called FabLabs (Fabrication Laboratories) have been held and promoted by the Center for Bits and Atoms (CBA) of the Massachusetts Technologic Institute (MIT).

FabLabs are equipped with a series of machines controlled by computer "to build (almost) everything": printers in three dimensions, cutter laser, numerical control (CNC) milling machines and an electronic laboratory. The global computation of FabLabs is led by the United States which boasts more than thirty, followed by the Netherlands (with 9), France (with 8), Spain (with 7) and Germany (with 6).5. Difficulties, development and assembly of a RepRap machine.A common problem of the RepRap printers is the necessity of an external computer connected to the printer, usually by USB (Universal Series Bus), with the appropriate device controllers. This computer does the following functions:It processes files with the three-dimensional definition of the objects to print (usually STL files) and transmits orders to the printer (using the so-called GCODE commands). The computer must remain connected during the printing.

The computer must remain connected during the printing. It serves to load and modify the printer software, for example to calibrate it after a material change.In order to avoid this dependence, there have been several variants of controllers that include memory card readers (normally SD cards), where you can burn files of GCODE orders from other computers and send them to the printer. In this case, it isn’t necessary to have external equipment, except for calibration.In order to be able to identify the problems in the assembly of a RepRap machine and be able to propose alternatives that improve the process, we decided to build a RepRap Mendel Prusa, but with a number of modifications.

This new machine was called RepRap Azara (Figure 8), following the custom of the RepRap movement in which areused names of personalities who contributed to the evolution study (in memory of Felix de Azara who was a Spanish naturalist of the 18th century).Figure 8: RepRap Azara.Firstly, the controller was replaced by one much more powerful, with enough capacity to perform all the work that is currently done with the external computer. In this way, the necessity of an external computer connected to the printer has been completely removed.For communication with the RepRap, an HTTP server is used (Hypertext Transfer Protocol, usually know as web server), installed on the new controller.

In this way, the printer use and the calibration is done by web, so it is possible to use the machine with smartphones, tablets, etc.In addition, it has made possible that the printer is able to transform three-dimensional designs in manufacturing orders, i.e., is able to transform STL files into GCODE instructions without external assistance.

This prevents that users have installed special applications, properly calibrated for the printer.Basically, the new RepRap variant replaces the electronics of the printer, replacing the original 8-bit controller (Arduino or similar) with a 32-bit microcomputer (Beaglebone) [13].This controller change has involved the development of:An expansion plate of Beaglebone, with capacity to control stepper motors, heaters, end stop and temperature sensors.A new resident software driver (firmware), which is capable of transforming the GCODE commands into motor movements, control temperatures, etc.A web server, allowing calibration, the sending of three-dimensional design (STL) files or GCODE files, direct control of motors and temperature, etc.The new design has also added advantages which are linked through a web server using:The printer can be used directly from any computer connected to the local network.The printer may be used at a distance, by configuring the router that connects it to the internet.A remote printer management can be done, which facilitates the maintenance of the machine.The website developed for the printer control allows its use from any computer which is able to visualize the models and send files through an HTLM form.The printers with SD card readers only print independently.

They still need an external computer to calibrate or modify the firmware, and it can’t be administered remotely.In addition, we have improved some aspects of the printer structure. We detected recurring problems in parts that allow the movement transmission of the z-axis motors to the threaded bars, called z-axis engines. The original parts, printed in PLA by a RepRap, don’t support an intensive use. In order to solve the problem, we replaced these parts with others of aluminum which have perfectly fulfilled their function.We also verified that the extruder blocked after using several hours because the melted plastic was too fluid and ascended by capillary action from the fuser to the extruder where the plastic fiber solidified and blocked. In order to solve the problem, we modified the design of the extruder and we incorporated an electric fan which can cool the upper fuser area, avoiding the melted plastic to ascend.But this fan cools the hot zone of the fuser.

This situation causes difficulty to get the temperature of plasticization and increase the energy consumption. In order to avoid this, we designed a piece adapted to the fan which diverts the air flow to the low area of the fuser. New parts have been manufactured with our own RepRap printer.6. Improvements and possible application fields.After analyzing the characteristics of RepRap printers and the complete assembly of one of them, we could suggest, in general, some changes of improvements:Edit the thermistors tables from the browser, and access to the complete skeinforge calibration.Add an user control to the web page. If you log in as an administrator you can manage print permissions, change of tail order, calibration, etc.Create a remote administrator of the machine which access a log to be able to monitor the printed parts, changes, problems, etc.

This remote administrator may be used to diagnose remote problems suffered by the printer.Add a video camera that could be used as a complement to the remote administrator, or to visualize at a distance the progress in the printing of the pieces. The images would be transmitted by the website of the printer. Improve the algorithm that makes movements in the printer to achieve more appropriate speeds and accelerations.Vary the design of the RepRap MendelMax, made of extruded aluminium profiles.Add an emergency button which instantly cuts the movement of the machine.  Modify the expansion plate to allow a touch screen connection. In this way, you can completely manage the printer without having to connect to its website.  Develop simple visualization software to facilitate the use of this type ofmachines to users without technical knowledge.

 Develop simple visualization software to facilitate the use of this type ofmachines to users without technical knowledge. It could also be considered in future works to improve the machines with new drives based on linear guides.Due to the self-replication potential of the machines, it is possible to cheaply distribute RepRap machines to individuals and communities (you can buy a complete kit of a RepRap printer by approximately $700), allowing them to create (or download from internet) products and complex objects without the need of expensive industrial machines. Application fields where RepRap machines could be used are: architecture, industrial design, engineering, robotic, education, toys, jewelry, medicine, feeding, etc, but it is always necessary to take into account part of the problems discussed above when we are developing new projects.7.

Conclusions.The RepRap technology will allow many small companies and engineering or architecture studies to have their own rapid prototyping machine.The low-cost, self-replicating character, GPL license, ease of assembly and existence of user and exchange communities are some of the features of RepRap machines that are valued very positively by users. The main problem is that RepRap printers have much less precision than other 3D printers.The RepRap machines still have a long way to go on improvement and research fields. That is the reason why the people working in this area are continuously increasing during the last years.In terms of social impact, this technology opens a wide range of possibilities for anyone can manufacture its spare parts without having to resort a technical service, or in the case of discontinued equipments, when it is not feasible to find that piece in the market.REFERENCES[1] Jones R., Haufe P., Sells E., Iravani P.

ditive Manufacturing without Layers: A New Solid Freeform Fabrication Process based on CNC Accumulation Yong Chen*, Chi Zhou, Jingyuan Lao Epstein Department of Industrial and Systems Engineering University of Southern California, Los Angeles, CA 90089Corresponding author: yongchen@usc.edu, (213) 740-7829ABSTRACTMost current additive manufacturing processes are layer-based, that is building a physical model layer-by-layer. By converting 3-dimensional geometry into 2-dimensional contours, the layer-based approach can dramatically simplify the process planning steps. However, there are also drawbacks associated with the layer-based approach such as inconsistent material properties between various directions. In a recent NSF workshop on additive manufacturing, it is suggested to investigate alternative non-layer based approaches. In this paper, we present an additive manufacturing process without planar layers.

In this paper, we present an additive manufacturing process without planar layers. In the developed testbed, an additive tool based on a fiber optics cable and a UV-LED has been developed. By merging such tools inside a liquid resin tank, we demonstrate its capability of building various 2D and 3D structures. The technical challenges related to the development of such a process are discussed. Some potential applications including part repairing and building around inserts have also been demonstrated.KEYWORDS: Additive manufacturing, 5-axis SFF, CNC accumulation, build around inserts.1. INTRODUCTIONCurrently most additive manufacturing (AM) processes are layer-based, that is building a physical model layer-by-layer. For example, as shown in Figure 1, to build a beam AB, a set of 2-dimensional layers are sliced and stacked together.

For example, as shown in Figure 1, to build a beam AB, a set of 2-dimensional layers are sliced and stacked together. By converting 3-dimensional (3D) geometry into 2-dimensional (2D) contours, the layer-based approach can dramatically simplify the process planning steps. However, there are also drawbacks associated with the layer-based approach such as inconsistent material properties between Z axis and other directions. In a recent NSF workshop on additive manufacturing (Bourell, et al., 2009), it was suggested to investigate alternative non-layer based additive manufacturing approaches. For example, to build the beam AB as shown in Figure 1, we believe a more ideal way is to move an accumulation tool along the direction AB. Hence, the building process would be much faster; more importantly, the material property of the built beams would be improved.Figure 1.

Hence, the building process would be much faster; more importantly, the material property of the built beams would be improved.Figure 1. An illustration of additive manufacturing process with and without layers.Another drawback of layer-based AM is its limitation on allowable motions between tools and work pieces. That is, by converting 3D geometry into 2D layers, most AM systems will only allow translational motions in X, Y, and Z axes. Consequently it will bring problems in building parts withembedded components (i.e. building around inserts). For example, for a popular AM process such as Stereolithography Apparatus (SLA), Kataria and Rosen (2001) identified some major problems in building around inserts such as laser shadowing and vat recoating. Both are mainly caused by the layer- based building processes. In comparison, multi-axis CNC machining, even with its great complexity in tool path planning, is getting more and more widely used in various industries.

In comparison, multi-axis CNC machining, even with its great complexity in tool path planning, is getting more and more widely used in various industries. Due to its flexibility in tool’s motion, multi-axis CNC machining such as 5-axis CNC milling, can significantly reduce building time, and improve building quality such as accuracy and surface finish (Apro 2009). Due to the well- known Moore’s law, computing performance per unit cost has been improved exponentially over the last thirty years. High performance personal computers (PCs) have enabled significant advances in the tool path planning of multi-axis CNC machining.In this paper, we present a novel additive manufacturing process that is not layer-based. Theprocess is named multi-axis CNC accumulation since it has great similarity to multi-axis CNC machining. As shown in Figure 2, CNC machining uses a machining tool to remove material that is in touch with the tool.

As shown in Figure 2, CNC machining uses a machining tool to remove material that is in touch with the tool. Hence for a given work piece (W) and tool path (Si) with tool orientation in each cutter location (Oj), . In comparison, the CNC accumulation uses an the constructed shape (M) will beM W =− ∪T + ( )S Oijaccumulation tool to add material that it touches. Hence the constructed shape will beM= ∪T + ( )S OiWork Piece (W)Machining Tools (T)CNC MachiningTool Path (Si + Oj)M W =− ∪T + ( )S OiLiquid ResinAccumulation Tools (T)Tool Path (Si + Oj)CNC AccumulationM= ∪T + ( )S OiFigure 2. A comparison between CNC machining and CNC accumulation.The paper is organized as follows. Section 2 describes the CNC accumulation process. Both its hardware and software modules are discussed. Section 3 presents our curing study of UV-curable resin.

Section 2 describes the CNC accumulation process. Both its hardware and software modules are discussed. Section 3 presents our curing study of UV-curable resin. Section 4 and 5 present our study on the attaching forces between the cured resin, base surface, and the curing tool. Section 6 discusses case studies of building various geometries by using our process. Section 7 presents two applications that we believe the process shows promises. Finally Section 8 concludes the paper.2. PROCESS DESCRIPTIONWe consider an additive manufacturing process based on ultraviolet (UV) curable liquid resin such as SLA. It is well known that UV light source can provide energy to convert liquid resin into solid parts. However, different from SLA, the curing tool in our process is kept merging under resin; hence the tool is capable of curing resin in various orientations.

However, different from SLA, the curing tool in our process is kept merging under resin; hence the tool is capable of curing resin in various orientations. A related problem by allowing the tool to directly contact with resin is that the newly cured resin may be attached to the curing tool instead of the base or previously built part (refer to Figure 3). We address the problem by studying the attaching forces related to the cured portion, which is always constrained between the base or the previously built part and the curing tool. It is found that the process can always be successful if: (1) certain type of coatings (e.g. Teflon film) are added on the tip of the curing tool such that the attaching force between the resin and the.tool can be reduced; (2) adequate over curing is ensured such that the attaching force between the newly cured resin and the base or the previously built part will be sufficiently big.

Hence the newly cured portion will always attached to the base or the previously built part instead of the curing tool. Our experimental results and related mathematical analysis are presented in Sections 3~5.Figure 3. The structure of the layerless under resin process.2.1. Hardware ComponentsIn the CNC accumulation process, it is desired to have an accumulation tool that can be easily translated and rotated. After considering various technologies including UV lamp, UV laser diode, and UV-LED, we selected a high power UV-LED from Nichia as the curing light source. A UV-LED is an inexpensive electronic light source with good properties such as low energy consumption, long lifetime and faster switching. The light is further focused with a sapphire ball lenses and transmitted through a quartz fiber optic light guide, both from Edmunds Optics. One side of the light guide, after applying a thin Teflon film, was merged in resin as the curing tool.

One side of the light guide, after applying a thin Teflon film, was merged in resin as the curing tool. A 5-axis table including the X, Y, Z axis translation and A, B axis rotation is used to provide desired motions. Hence the light guide can be translated and rotated to provide UV light in various orientations. The hardware setup of our CNC accumulation system is shown in Figure 4. In addition to the optical and motion execution devices, it also includes a force sensor and related controller to provide feedbacks.Figure 4. The hardware system of the CNC accumulation process.Figure 5. The software system of the CNC accumulation process.2.2.

The hardware system of the CNC accumulation process.Figure 5. The software system of the CNC accumulation process.2.2. Software ComponentsFor a given geometry, a popular CNC machining software system, Mach3 CNC control application (www.machsupport.com) is used to convert the shape file of the geometry into numeric control G-code.As shown in Figure 5, a CNC accumulation controller was developed to read in the G-code and accordingly send commands to a KMotion control board, which can drive up to 8 axis motions and control more than 20 digital IOs. Hence a desired model can be built by controlling the LED light (on and off) and its 5-axis motion.3. CURING OF RESINAs discussed before, the cured resin in our process is always constrained between the curing tool and the previously built part. Hence, a proper gap has to be maintained for such a constrained curing process.

Hence, a proper gap has to be maintained for such a constrained curing process. To ensure the newly cured portion will always attach to the previously built part, the attaching force between them must be bigger than the one between the newly cured portion and the curing tool. Therefore, if the gap is large, longer curing time or slower scanning speed is required in order to achieve sufficient attaching force between the cured portion and the previously built part. However, too long curing time will also increase the building time and enlarge the sizes of the cured portion, which leads to a lower resolution of the built part. Hence, in this section, we first discuss the relations between the cured resin including its shape and sizes related to the provided light energy (i.e. exposure time and scanning speed). Such understanding will provide the basis for the attaching force study as discussed in Sections 4 and 5.3.1.

exposure time and scanning speed). Such understanding will provide the basis for the attaching force study as discussed in Sections 4 and 5.3.1. Principle of Resin CuringThe curing process for a laser beam has been extensively studied (Jacobs 1992). According to the Beer- Lambert exponential law of absorption, the laser exposure will decrease exponentially with depth z. The dependence of curing depth Cd upon the maximum exposure at the resin surface follows the equation:(3.1) In our process, the focused spot of the UV-LED has a Gaussian distribution, which is quite similar to a laser beam. Hence we assume the curing behaviors are the same for these two kinds of light sources. As extensively studied in the SLA process (Jacobs 1992), a critical energy exposure threshold (Ec), can be found for a given type of liquid resin. When energy input per unit area is less than the minimum energy requirement of Ec, the material will remain as a liquid or gel.

When energy input per unit area is less than the minimum energy requirement of Ec, the material will remain as a liquid or gel. The material will then be removed during the post-processing processes due to the lack of mechanical strength. Curing models also indicate that the input energy determines the depth of penetration of resin (Dp) which should be bigger than the layer thickness.C d=Dln(E/E c)pmaxWhen the LED spot scans in a straight line with a constant velocity over the resin surface, the X-Y plane is coincident with the resin surface, the X axis is corresponding to the scanning direction and the Y axis related to the curing width, while the Z axis is perpendicular to the resin surface and also related to the curing depth. Setting y=y* and z=z* when E(y,z) = Ec, according to the parabolic cylinder analysis (refer to Figure 6), the curing width and the curing depth satisfy the following parabolic equation of the cured line’s cross-section:Figure 6.

Parabolic cylinder of a cured line (Jacobs, 1992).ay2*+bz= (3.2)cwherea=2 2 W 0,b=1 Dpandc=ln⎧ ⎪ ⎨ ⎪ ⎩2 πP L W V E 0 s c⎫ ⎪ ⎬ ⎪ ⎭,W D V P E are all constants.0 ,,,,psLcFurther deduction of the mathematical analysis can obtain the following important equation:L w=BC d D 2p(3.3)Where wL is the line width, B is the light spot diameter. This equation shows that the cured line width is proportional to the laser diameter and also proportional to the square root of the ratio between the curing depth and the resin penetration depth.3.2. Experiments and AnalysisTo verify the curing behavior of a focused UV-LED used in our system, we performed experiments based on point curing. As shown in Figure 7.(a), the point curing process is conducted by turning on the UV- LED for a certain time. The cured resin is then detached and measured to study the relation between curing shape and curing energy.

The cured resin is then detached and measured to study the relation between curing shape and curing energy. The sizes of the cured bullet were measured under different exposure time. The curve between the input energy (measured by the curing time) and the curing depth are shown in Figure 7.(b). The curve between curing depth Cd and ln(E) matches equation (3.1) quite well, which is established for the SLA process.Figure 7. The curing depth versus the exposure energy for point curing.Further, we can approximate the relation between the curing depth Cd and the curing time T for theresin used in our system as: =dCT 1.0513 ln( ) 1.2537∗+. (3.4). In this work, and the relation between the curing width Lw and the curing depth Cd as we further assume the constant number between the width and depth for point curing and line curing are the same.

In this work, and the relation between the curing width Lw and the curing depth Cd as we further assume the constant number between the width and depth for point curing and line curing are the same. Thus, the curing depth and line width under different exposure time or scanning speed can be approximately computed.L w=1.5016Cd4. ATTACHING OF POINT CURED RESINAs shown in Figure 3, the newly cured portion of resin in the CNC accumulation process is attached to the curing tool as well as the basis or the previously cured part. We model the attaching forces related to the two interfaces of the newly cured resin to ensure the building process can be successful.4.1. Attaching Force ModelingDenote F as a force over an attaching surface area A. Suppose FTool is the attaching force between cured resin and the tool; and FBase is the attaching force between cured resin and the basis or the previously cured part.

Suppose FTool is the attaching force between cured resin and the tool; and FBase is the attaching force between cured resin and the basis or the previously cured part. In order to successfully separate the cured portion of resin from the curing tool, .FBaseThe attaching forces are mainly determined by two factors: (1) The types of materials that are in contact with the cured resin. Suppose α is the coefficient of attaching forces related to the type of material (i.e. attaching force over a unit area). Hence it is desired to haveToolα as small as possible whileBaseα as big as possible;(2) The attaching area A. As discussed in Section 3, the area A can be estimated based on theinput energy.Hence to ensureFBaseF>Tool, we know:αBaseiABaseα>TooliA Toolorα αBaseTool>A Tool ABase.F>ToolFigure 8.

Attaching force study for point curing.As shown in Figure 8.(a), after the UV-LED has been turned on for a certain time TCuring, portion of resin between the LED and the base surface will be cured. We then turned off the UV-LED and slowly move it upwards. The cured resin will attached to either the base if . Based on the study in Section 3, we know the surface area A1 is fixed for a constant curing time. However, for different gap value d, the cross section area A2 will be different. Further A2 will decrease when d increases.F 2F> 1, or the LED ifF 1F> 2Hence for a given TCuring, we can perform a set of experiments by slightly varying d values to identify such a critical state dC where , the cured resin will attached to the base. As shown in Figure 8.(b), similar experiments can be performed for different curing time to identify a detachable region for the two types of materials that are used in the tool and the base (i.e. Teflon film and cured resin in the figure).

Teflon film and cured resin in the figure). All the red points in the figure correspond to the critical state for the specific curing time. The shaded portion under the curve denotes the detachable region, i.e, all the settings in this region can guarantee the successful separation of the cured resin. Further, can. Accordingly we know ifF 1F= 2dd<Cbe computed since the ratioA 1 A 2related to dC can be estimated.(a) The comparison between the different bases. (b) The comparison between different tool coatings.Figure 9. The relation between the gap and curing time for point curing.4.2. Experiments and AnalysisBased on a similar approach, various types of materials have been studied. Figure 9.(a) shows the comparison between Teflon film and different base materials including plastic, glass and cured resin. It shows that:α 2 α 1(1) For all the base materials, the exposure time should be increased for an increased gap distance.

It shows that:α 2 α 1(1) For all the base materials, the exposure time should be increased for an increased gap distance. (2) For the same exposure time, the cured resin as the base can have the largest gap, i.e, the systemwith the cured resin as the base has the largest safe region.Similar experiments have been performed on the curing tool side. By using cured resin as the base, different types of coatings including Teflon film, plastic film and fiber optic head (i.e. without any coating) are tested. The results are shown in Figure 9.(b), which illustrates that:(1) The Teflon film as the intermediate material gives us the largest safe region. (2) The fiber optic head has the smallest safe region. Hence the process will be more likely to fail ifno coating is applied on the tip of the tool.To verify the experimental result, we also took two arbitrary points for different curing time, one locating inside the safe region and one outside the safe region.

We performed the aforementioned experiments for the sampling points. The test results are in accordance to the conclusion we draw from the figures.areaBased on the mathematical analysis and experimental result in Section 3, the ratio of the attaching A 1 A 2can be approximated for the critical gap dC. Sincefor the equilibrium states,F 1F≈ 2α 2 α 1≈Hence the force coefficient ratio to the curing time can be computed, which is shown in Figure 10.αCoating αBaseαCoating αBaseFiber optic Plastic Film(a) The comparison between the different bases. (b) The comparison between different tool coatings.Figure 10. The relation between the force coefficient ratio and curing time for point curing.From the figure, it can be observed that:(1) For all the materials, the attaching force coefficient ratio is decreasing as the curing time is increasing.

That is, for a longer curing time, the attaching force on the coating is increasing faster than the attaching force on the base.V(2) Among all the coating and base materials, the pair that would give us the best performance is by using cured resin as the base and Teflon film as the tool coating. The average force ratio between them is around 0.4.5. ATTACHING OF LINE CURED RESINFigure 11. Attaching force study for line curing.Similar studies have been performed for line curing, in which the UV-LED is constantly moving in a direction that is orthogonal to the curing direction.A 1 A 2.5.1. Attaching Force ModelingDuring the line curing, the moving direction of the curing tool is different from its curing direction. Hence the attaching forces are mainly shearing forces as shown in Figure 11. We can use FTool as the shearing force between cured resin and the tool, and FBase as the shearing force between cured resin and the basis or the previously cured part.

We can use FTool as the shearing force between cured resin and the tool, and FBase as the shearing force between cured resin and the basis or the previously cured part. Hence in order for the cured portion of resin to successfully separate from the curing tool, . Using the same procedure as discussed in Section 4, we can establish the relation between the gap distance and the scanning velocity. As shown in our test results, the combination of resin base and Teflon film also gives us the best performance for line curing.FF>BaseTool5.2. Experiments and AnalysisBy moving the UV-LED at a constant speed V, portion of resin between the LED and the base surface will be cured. The cured resin will attached to either the base if . Based on the study in Section 3, we know the surface area A1 is fixed for a constant speed. The cross section area A2 will decrease for a bigger gap d.

Based on the study in Section 3, we know the surface area A1 is fixed for a constant speed. The cross section area A2 will decrease for a bigger gap d. Hence for a given V, we can perform a set of experiments by slightly varying d values to identify such a critical state dC where . Accordingly a critical point for two types of materials that are used in the tool and the base can be identified. The experiments are repeated for different speeds. Hence a detachable region can be generated, in which all the settings can guarantee the successful separation of the cured resin. Figure 12.(a) shows the comparison between Teflon film and different base materials including plastic, glass and cured resin. Figure 12.(b) shows the comparison between cured resin and different coatings including Teflon film, plastic film and fiber optic head.F 2F> 1, or the LED ifF 1F> 2F 1F= 2(a) The comparison between the different bases. (b) The comparison between different coatings.Figure 12.

(b) The comparison between different coatings.Figure 12. The relation between the gap and scanning velocity for line curing.αCoating αBaseαCoating αBase(a) The comparison between the different bases. (b) The comparison between different coatings.Figure 13. The relation between the force coefficient ratio and curing time for line curing.The relation between the force coefficient and the speed has also been studied. The relations areshown in Figure 13. Compared to the point curing result as shown in Figure 10, it can be seen that:(1) The force coefficient ratio for line curing is consistent to that of point curing. The relative sizes and are:ofvarioustestedmaterials>α>α>ααPlastic_ ResinCuredGlass for both point and line curing.Teflon Film _αα>α>α>Plastic Film _Fiber Optic __ ResinCuredTeflon Film _(2) The force coefficient ratio for line curing is closer to a constant.

In addition, such ratio valuesare close to the average values ofαCoating αBasein point curing.6. CASE STUDIES OF VARIOUS GEOMETRIES To verify the proposed CNC accumulation process, a set of case studies including 2D and 3D geometries have been carried out. The experimental results verified that the presented method can build different types of geometries. In addition, it can achieve material accumulation in different directions. In all the case studies, Teflon film was used as the coating on the fiber optic head.6.1. 2D Curves As shown in Figure 14.(b), a simple square frame was first tested to verify the feasibility of our process. The tool paths we used in the test are illustrated in Figure 14.(a). Both point and line curing as discussed in Sections 4 and 5 are tested in the case study. A more complex 2D shape, a pattern of “I ♥ LA”, was used to test the linkage between Mach3 CNC application and our controller.

A more complex 2D shape, a pattern of “I ♥ LA”, was used to test the linkage between Mach3 CNC application and our controller. The pattern was defined in a geometric file and converted into NC G-code. Following the control commands, the part was successfully built and shown in Figure 14(c). All the parts in the test cases were built on resin base.(a) Tool paths (b) Squre frame (c) “I love LA” character strokeFigure 14. Results of 2D curve study.Figure 15. Test of a Figure 16. Results of 2D surface study: adding characters2D planar surface. and dragon textures on a plastic sissor plate.6.2. 2D SurfaceTest cases on building 2D surfaces were conducted. In order to build a part that is above the bottom of the tank, supports have to be added. We used the same approach as shown in 2D curves to build such supports. A 2D surface can be built by repeating scanning a set of 2D curves.

We used the same approach as shown in 2D curves to build such supports. A 2D surface can be built by repeating scanning a set of 2D curves. A certain angle of the tool was used to keep the newly cured resin bonding with the already cured part. An example of the finished 2D planar surface is shown in Figure 15, which verifies the feasibility of our process on building 2D surface. After the curing process, it is noticed that some bending occurred due to the shrinkage of cured resin. Reducing such deformation can be learned from the SLA process and is a research topic in our future work.Besides cured resin, as discussed in Sections 4 and 5, other types of bases can also be used in the CNC accumulation process. As shown in Figure 16, an acrylic plastic plate is used as the base. Some designed characters and decorating patterns can be added on top of it.6.3. 3D CurveOne of the main advantages of the CNC accumulation process is that no layers are needed when building 3D parts.

3D CurveOne of the main advantages of the CNC accumulation process is that no layers are needed when building 3D parts. Hence it can dramatically improve the manufacturing efficiency and achieve the material consistency in desired directions. A 3D half circle in XZ direction was built as shown in Figure 17. Supports are added to facilitate the building process. In this process, since the rotation table is used to keep the curing tool pointing to the tangent direction of the circle, the built arc can be smooth with layer- related artifacts.A more sophisticated case on spatial characters has also been performed. In this test, we try to add some characters “USC FIGHT ON” on the inner surface of an existing plastic bowl. We first computed the G-Code for the characters. We then computed the rotation of the bowl and the rotation of the curing tool related to the shape and position of each character. The part was successfully built on the plastic bowl.

We then computed the rotation of the bowl and the rotation of the curing tool related to the shape and position of each character. The part was successfully built on the plastic bowl. The bonding between the characters and the bowl was also verified to be strong.Figure 17. Test of a 3D curve. Figure 18. Test of spatial characters.Figure 19. Test of a 3D cylindrical surface.6.4. 3D SurfaceSimilar to a 2D surface (refer to Figure 15), we tested a 3D surface (a 3D cylindrical surface) to demonstrate the feasibility of our process. Some supports were added to facilitate the building process. The built result is shown in Figure 19. Even though the building process was finished successfully, the part quality of the built surface needs to be improved in the future.Discussion: As illustrated in Figure 1, the CNC accumulation process can overcome the anisotropic behavior of 2D and 3D curves fabricated by layer-based manufacturing processes.

However, 2D and 3D surfaces that are made by the CNC accumulation process will still have anisotropic behavior (refer to Figure 15 and 19). Compared to the layer-based manufacturing processes, the CNC accumulation process can have more freedom in its tool path planning. Hence the build orientation can be adjusted to have an anisotropic behavior that fits design requirements better. In addition, neighboring layers in CNC accumulation can have a better bonding due to the rotation of the curing tools.7. APPLICATIONS7.1. Part Repairing and ModificationOur process has similarity to Laser Engineered Net Shaping (LENS), Direct Metal Deposition (DMD), and laser cladding, which have been widely used in repairing metal parts and molds [5~7]. However, we did not find a SFF process that can repair a plastic part or mold.

However, we did not find a SFF process that can repair a plastic part or mold. For most rapid prototype processes, even though it is quite flexible for them to directly build a new part, it is far less flexible for them to repair or modify an existing part. We believe the presented CNC accumulation process can fill such gaps. Our system can repair a plastic part, and also enable desired features to be added on a model that has been built by using other SFF processes such as SLA and SLS. A test case to demonstrate the effectiveness of the proposed process is shown in Figure 20.We first use SLA to built a pig model. Assuming a new feature in the front of its chest is later identified and required to be added. Instead of throwing away the existing part and rebuilding the modified CAD model, we convert the new feature into related tool paths. Then the SLA part was merged under resin and fixed in the tank.

Instead of throwing away the existing part and rebuilding the modified CAD model, we convert the new feature into related tool paths. Then the SLA part was merged under resin and fixed in the tank. After necessary calibration and positioning, the new feature was built automatically based on the given tool path. After the building process, the modified part was take out and post-processed. The bonding between the newly added feature and the SLA model was tested. It strength was found to be satisfactory.SLA modelPlan repair pathTool PathsFigure 20. Test of part repairing and modification.7.2. Building around InsertsTraditional rapid prototype process can build very complex part. However, most of them such as SLA and SLS can only build part with one type of material. Even though some of the processes can combine different kinds of materials such as Objet’s Connex printer, the materials, however, have similarAdded featureproperties.

Even though some of the processes can combine different kinds of materials such as Objet’s Connex printer, the materials, however, have similarAdded featureproperties. In most products design, various types of materials are required due to different functional requirements. To address the problem, the idea of “building around inserts” has been presented before, in which existing components made by other manufacturing processes can be integrated in the SFF processes.

To address the problem, the idea of “building around inserts” has been presented before, in which existing components made by other manufacturing processes can be integrated in the SFF processes. We believe the presented CNC accumulation process can potentially address most requirements of “building around inserts” due to its flexibility in building parts in varying directions and also on different base materials.Several test cases including a multi-functional scissor, a decoration part with embedded opticalfibers and a model for resistor wrapping are used to verify our idea.(1) Adding a spring in a multi-functional scissor: A simple scissor was built based on two pieces of laser-cut acrylic plastic plates. In this test, we would like to add a spring to connect the two pieces of the body.

In this test, we would like to add a spring to connect the two pieces of the body. Even though SLA or SLS can build a plastic spring without any problem, it is desired to have a metal spring since the spring is required to undergo comparatively large compressing and stretching forces. We designed tool paths and used our system to connect the spring with the two plastic pieces. The built part with the spring is shown in Figure 21.Figure 21. Test of adding a spring. Figure 22. Test of adding a magnet bar.(2) Adding a magnet bar in a multi-functional scissor: In this test, we would like to add a magnet bar on the surface of a plastic plate such that it can hold some tiny metal parts. The tool path for connecting the magnet bar on the plastic plate of the scissor is shown in Figure 22.

The tool path for connecting the magnet bar on the plastic plate of the scissor is shown in Figure 22. The built scissor with both the spring and the magnet bar for functional testing is also shown in the figure.(3) Building a part to connect multiple optical fibers: Optical fibers are inexpensive with good properties. However, their fixture in existing parts is troublesome due to their small sizes (<0.25mm). In this test, we would like to build a decoration part by connecting 10 single optical fibers into a star shape. We designed a fixture and use our system to build a photopolymer part in start shape (refer to Figure 23). After the building process, the optical fibers have been embedded in the part. Hence, interesting lighting patterns can be achieved by controlling the light source provided at the other end of the optical fibers.Figure 23.

Hence, interesting lighting patterns can be achieved by controlling the light source provided at the other end of the optical fibers.Figure 23. Test of building a decorating part with optical fibers.(4) Wrapping a resistor: In this test, we show the capability of our process to wrap existing electric components. Two resistors were used to be integrated in a built part. The tool path and related built results are shown in Figure 24.Figure 24. Test of wrapping resistors.8. CONCLUSIONA new solid freeform fabrication process based on CNC accumulation was presented in this paper. The proposed process builds parts under liquid resin and can build parts without planar layers. Five axis motion controls are incorporated in the process to achieve desired movements between accumulation tools and built parts. The curing mechanism of our process has been studied, which is found to be similar to the SLA process. Different attaching forces of cured resin have also been studied.

The curing mechanism of our process has been studied, which is found to be similar to the SLA process. Different attaching forces of cured resin have also been studied. The results show that a combination of cured resin as base and Teflon film as the surface coating can provide the best performance for the separation of cured resin from the tool. An integrated system with various hardware and software components has been built. Case studies of various geometries have been performed. The experimental results demonstrate the effectiveness of the proposed method for applications such as (1) plastic part repairing, (2) plastic part modification, and (3) building around inserts.We presented a proof-of-concept system for a novel AM process. There are plenty work that can be done to improve the developed system and the related process. Some current work we are investigating includes: (1) developing curing tools that have a much smaller size (e.g.

Some current work we are investigating includes: (1) developing curing tools that have a much smaller size (e.g. 0.25mm); (2) incorporating collision detection between the built part, inserted part, and the tool; (3) integrating sophisticated path planning algorithms into the process planning; (4) developing new applications that are enabled by our process.REFERENCES1. Apro, K. Secrets of 5-Axis Machining, Industrial Press Inc., New York, NY, 2009. 2.Bourell, D., M. Leu, and D. Rosen, NSF Workshop - Roadmap for Additive Manufacturing: Identifying the Future of Freeform Processing, Washington, D.C., March 30-31, 2009. Jacobs, F. P. Rapid Prototyping and Manufacturing: Fundamentals of Stereolithography, Society of Manufacturing Engineers, Dearborn, MI, 1992.3.4. Kataria, A. and D. W. Rosen. “Building Around Inserts: Method for Fabricating Complex Devices in Stereolithography”. Rapid Prototyping Journal. Vol. 7, No. 5, pp. 253-261, 2001.5. Mudge, R. P. and N. R.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceADDITIVE OS: AN OPEN-SOURCE PLATFORM FOR ADDITIVE MANUFACTURING DATA MANAGEMENT & IP PROTECTIONEvan P. Diewald*Department of Mechanical Engineering, Carnegie Mellon University, Pittsburgh, PA 15213AbstractThe additive manufacturing (AM) digital thread presents unique challenges for data management and security. While proprietary software packages solve many issues, they can be expensive and lacking in customization. Additive OS is an open-source platform for importing, sharing, organizing, and querying AM data. Man-in-the-middle attacks, secure print licensing, and IP theft are addressed using custom smart contracts, ontology is preserved with a NoSQL database and directed acyclic graph (DAG) representations, and peer-to-peer content delivery facilitates low- latency file transfer.

The application includes a browser-based graphical user interface, but developers can access the underlying API to invoke sophisticated queries, add functionality, or run the lightweight client on low-resource hardware.IntroductionCybersecurity Threats in AMOne of the features that distinguishes additive manufacturing (AM) from other forms of manufacturing is its virtualizable workflow. In this digital thread, the fabrication cycle is accompanied by a digital twin, which takes the form of design files, build planning parameters, process monitoring data, invoices, and other media. By nature, the process is highly iterative, as part geometries and machine inputs may be optimized in response to simulations or mechanical testing. Given that 3D printed components are particularly common in mission critical applications such as aerospace and defense, it is essential that regulators can confidently trace a part’s entire lifecycle.

However, the digital thread is vulnerable to a number of attack vectors that challenge conventional audit trails. Existing research has shown that saboteurs can relatively easily embed micro-defects in CAD files [1]–[3], alter machine input files and firmware [4], [5], or manipulate in situ monitoring data, which is increasingly used as a means of process qualification [6]. These subtle changes are non-trivial to detect but can have a drastic impact on a part’s mechanical properties. Cybersecurity is further confounded by the decentralization of the AM supply chain, which includes feedstock suppliers, engineering firms, and machine shops.As a rapid prototyping tool, AM is often used to fabricate sensitive or pre-production designs, making intellectual property (IP) theft a pressing concern for many users. In a survey of experts from industry, academia, and governmental organizations, Kurpjuweit, et. al.

In a survey of experts from industry, academia, and governmental organizations, Kurpjuweit, et. al. found that IP digital rights management and counterfeit prevention were considered top issues in AM cybersecurity, after threats related to sabotage and traceability [7]. One of the more eye-opening results from the survey was an anecdote that, when sharing confidential designs with an external firm, some designers prefer to mail a physical copy of the part rather than digitally share the CAD file. Such examples are indicative of the real-world challenges in establishing trust among distributed stakeholders in the AM supply chain. Typically, proving ownership of novel material254requires legal protection in the form of copyrights or patents.

Typically, proving ownership of novel material254requires legal protection in the form of copyrights or patents. In addition to the lengthy application process and significant legal expenses, this conventional approach often requires public disclosure of early-stage prototypes.AM and Blockchain“Blockchain”, or distributed ledger technology (DLT), is a broad term used to describe decentralized peer-to-peer networks of individual nodes that maintain an immutable, append-only record. Proposed state changes, or transactions, are bundled into the atomic units of the data structure, called blocks. Because each block contains the cryptographic hash of the previous block, the integrity of the ledger is validated by calculating each block’s hash in sequence. Blockchains vary in consensus protocols, which are the algorithms that provide an economic incentive for each individual node to keep an accurate and honest version of the current world state.

Blockchains vary in consensus protocols, which are the algorithms that provide an economic incentive for each individual node to keep an accurate and honest version of the current world state. “Proof of Work”, which was popularized by Nakamoto with the introduction of Bitcoin in 2009 [8], requires nodes to compete to “mine” new blocks of transactions by finding the solution to computationally-intense cryptographic puzzles1. The Ethereum blockchain is currently in transition from Proof of Work to a more efficient protocol known as Proof of Stake2. The upcoming launch of “ETH 2.0” aims to increase network capacity and reduce transaction fees without sacrificing security [9]. Other distinctions among DLT frameworks include the privacy of the network and the types of transactions that they support. Blockchains can typically be classified as private, meaning they have restricted access (e.g. Hyperledger Fabric), or public, meaning they are fully open (e.g. Bitcoin).

Blockchains can typically be classified as private, meaning they have restricted access (e.g. Hyperledger Fabric), or public, meaning they are fully open (e.g. Bitcoin). Consortium chains, maintained by several firms in the same industry, are a compromise between the two extremes. An important innovation established by Ethereum was the introduction of smart contracts, which are computer programs that interact with a blockchain through pre- defined functions and business logic. Each transaction with a smart contract is recorded on the shared ledger, providing a resilient audit trace.Cybersecurity experts in academia and industry recognize that blockchains can be applied to the AM supply chain to increase accountability. Research articles have presented theoretical frameworks illustrating how DLT could secure various aspects of the digital thread. In [10], the authors propose a private ledger to track the lifecycle of parts through a use case for the aerospace industry.

In [10], the authors propose a private ledger to track the lifecycle of parts through a use case for the aerospace industry. A conceptual approach for a blockchain-based AM data management and exchange system was outlined by Papakostas, et. al. [11]. Alkhader et. al. produced a smart contract with functions where designers and printers trace an example development sequence through a trustless interface3 [12]. A sample scenario was demonstrated by calling the functions within the Remix IDE. The “Secure Additive Manufacturing Platform” (SAMPL) defines a general smart contract in which part integrity and copyright protections could be validated from manufacturer to customer using RFID tags [13]. Kennedy, et. al. used lanthanide nanomaterial chemical signatures to print1 More specifically, miners attempt to find a “nonce” value that, when appended to the block’s existing data, results in a block hash with a certain number of leading zeros.

The number of leading zeros is known as the “difficulty”, and it is adjusted in response to the computing capacity of the network to ensure consistent block times. 2 In Proof of Stake, nodes “stake” a large amount of capital in exchange for the rewards associated with mining blocks. Dishonest or unreliable nodes forfeit all or a portion of their initial investment. 3 The smart contract required a third-party “Certificate Authority” to sign off on each transaction.255unique QR codes in the fused deposition modeling process. The codes could be scanned with a smart phone to log processing steps as Ethereum transactions [14].Commercial solutions are more complete, but they involve proprietary technology and costly licensing agreements. General Electric Additive filed a patent application in 2018 for a distributed ledger system that verifies that the build file and material lot have not been altered before reaching the machine [15].

General Electric Additive filed a patent application in 2018 for a distributed ledger system that verifies that the build file and material lot have not been altered before reaching the machine [15]. GE and Microsoft joined efforts on their “TRUEngine” platform, which uses a private fork of the Ethereum network to track components in aircraft engines throughout their lifecycle [16]. Start-ups like Cubichain Technologies have demonstrated a proof of concept to “protect the digital data stream for additively manufactured aerospace titanium parts” [17]. While the details of this confidential technology are limited at this time, Cubichain is built on the MultiChain private blockchain platform (based on Bitcoin Core), whose commercial license is currently priced at $5,000/year per node [18].At this time, the authors are not aware of an open-source application that facilitates AM data management in conjunction with smart contracts for independent traceability and IP protection.

Commercial platforms lean toward private architectures which are fast and secure but require more expensive infrastructure. Furthermore, as ultimately centralized solutions (nodes are maintained by a single organization), they do not take advantage of the distributed nature of public blockchains. By design, trusted networks cannot be independently audited by external parties, such as regulators. For large firms with vertically integrated AM supply chains, this may be acceptable, but many design shops and laboratories outsource powder supply, manufacturing, post-processing, and/or characterization work. One of the most significant benefits of 3D printing is the ability to manufacture at the point of need, which requires a reliable source of truth between geographically or organizationally disparate entities.

Additive OS balances privacy and performance concerns with the benefits of a fully trustless and objectively certifiable digital thread.AM Data ManagementUnder guidance from experts in industry, academia, and national standards institutions, America Makes published a Strategic Guide for AM Data Management and Schema in 2019 [19]. The committee voted “the need for unique, unified data identifiers for AM data” as one the most significant gaps related to AM data management. This convention would allow for more useful analytics, e.g. drawing “cross comparisons” between a given part’s process conditions, characterization data, and mechanical properties [20], [21]. AM facilities often store massive amounts of data in an ad hoc manner, limiting the usefulness of Big Data analytics and machine learning [22]. For users that opt for a unified database, they must choose between costly commercial packages, limited open-access platforms, or home-grown solutions.

For users that opt for a unified database, they must choose between costly commercial packages, limited open-access platforms, or home-grown solutions. GRANTA:MI [23] is a proprietary application for importing, storing, and querying AM data. Users choose between a selection of closed-source, pre-defined schemas that include process parameters, mechanical testing data, and URLs to images stored externally. Senvol Database [24] is a compilation of 3D printing machines and materials, meant to aid in selection of, for instance, polymer filament that will meet certain strength requirements. To access the underlying data and application programming interface (API), a license is required. Startups Link3D [25] and Authentise [26] integrate business operations, logistics, and AM workflow tracking within their proprietary manufacturing execution software.256In the academic sector, efforts are mainly focused on establishing an appropriate representation for AM data [27]–[31].

Computational ontologies formalize the inheritance structure of a dataset [32]; they have been successfully applied to domains like medicine and biology to extract meaningful relationships and maintain consistency [33]. This representation - allows data scientists to use advanced deep learning techniques such as graph neural networks, which utilize the adjacency between nodes as an intuitive feature set for regression and classification models [34], [35]. Establishing a common ontology makes the dataset easier to query, and new entries can be imported using existing templates.

Establishing a common ontology makes the dataset easier to query, and new entries can be imported using existing templates. Some databases and querying languages, including GraphDB [36] and SPARQL [37] are specifically designed to capture semantic structure, but ontologies can also be represented in general-purpose NoSQL platforms such as MongoDB [38], [39].In terms of web applications and platforms that integrate AM data ontology and workflow management, open-source solutions are significantly more limited than commercial competitors. The most polished application was developed by NIST [28], but new users must apply for special permission, and all data is completely public.

The most polished application was developed by NIST [28], but new users must apply for special permission, and all data is completely public. As highlighted in the America Makes guide, while sharing insights can enrich collective knowledge, security and privacy remain paramount in most cases.MethodsSmart ContractsAs opposed to their legal counterparts, smart contracts do not require third-party authentication to establish an immutable and objective audit trace. While smart contracts are most common in financial applications, such as decentralized exchanges [40], [41], they are also used to govern supply chains [42], manage healthcare records, and provide distributed voting mechanisms [43]. Most modern blockchains have support for smart contracts, but a comparison of developer platforms is beyond the scope of this article.

Most modern blockchains have support for smart contracts, but a comparison of developer platforms is beyond the scope of this article. Additive OS processes transactions on the Ethereum network [44], which interprets smart contracts written in Solidity, an object-oriented programming language that compiles high-level scripts into machine-readable bytecode [45]. Ethereum has relatively low transaction fees and adequate speeds, and it is by far the most popular blockchain for decentralized applications due to broad developer support and extensive tooling. Importantly, Additive OS decouples transaction signing from the application itself, meaning users do not have to hard-code their private keys into the system. Transactions are signed using Metamask, a browser extension-based Ethereum wallet with over 1M users [46].Additive OS deploys two smart contracts to address the most pressing AM cybersecurity threats of man-in-the-middle (MITM) attacks and secure design licensing.

The primary role of the AMProject smart contract is to provide methods for associating authenticated files with a project. Each Project object is defined by a unique index, its author’s ETH address, and an array of File objects corresponding to verified content. Each File contains the SHA3 checksum of an added file and the address of the user who signed the transaction. Thus, the legitimacy of a shared file can be confirmed by comparing its hash to what is stored in the smart contract (Figure 1). The inputs to one-way hashing algorithms like SHA3 cannot be reverse-engineered, ensuring that the257confidential contents of the file are not exposed to the public ledger [47]4. Even if Additive OS’s database cluster is compromised or deleted, the permanent contract and its associated transactions are not lost.

Even if Additive OS’s database cluster is compromised or deleted, the permanent contract and its associated transactions are not lost. Similar contracts have been demonstrated to secure the AM supply chain against tampering, but the applications were implemented in a toy setting and not integrated in a user interface [12]. In the first use case, Additive OS is used to expose a MITM attack on a turbine blade, similar to the dr0wned threat demonstrated by Belikovetsky, et. al. [1]. However, the same smart contract can be generalized to preserve the entire digital thread through substantiation of process monitoring data, invoices, and test results.Figure 1. Schematic of the AMProject smart contract, which protects against man-in-the-middle attacks.The second smart contract, AMLicense, addresses more subtle threats related to IP rights and print licensing.

A new license instance must be defined by the ETH wallet addresses of the licensor and licensee, the checksum of the part’s design file, and the number of licensed prints. Like the AMProject smart contract, by storing only the file’s hash, the sensitive file contents are not exposed to the public blockchain. Again, this can help protect against MITM attacks, but in this scenario, the intention is to establish ownership of the IP without public disclosure. If the licensee (e.g. a service bureau or fabrication shop) were to copy the design file for themselves or share it with a competitor, the smart contract transaction persists as a timestamped attestation of the original deal. As a trustless interaction, arbiters could verify the transaction contents without relying on the either party’s centralized and editable records.

As a trustless interaction, arbiters could verify the transaction contents without relying on the either party’s centralized and editable records. While scenarios like this were once a theoretical exercise, blockchains’ growing acceptance as a tamperproof audit trail is starting to hold real legal weight; several states have begun to pass legislation accepting smart contract transactions as admissible evidence [48]–[50].4 Moreover, concise checksums keep gas fees low.258When a print is executed by the licensee5, they log another transaction that includes an operator ID and the hash of the build report. This produces a link to a digital “certificate of authenticity” that can be carried with the physical part as a QR code or RFID tag. Further down the supply chain, the underlying transaction details can be independently audited by end users, and unauthorized duplicates (i.e. parts without a unique and valid certificate) can be easily identified.

parts without a unique and valid certificate) can be easily identified. The AMLicense smart contract (Error! Reference source not found.) demonstrates the benefits of using a public ledger like the Ethereum Mainnet, as opposed to private architectures implemented by some firms [16], [17]. By the time a 3D printed component is installed by a customer, its digital fingerprint may have crossed between several different organizations, from feedstock suppliers and designers to simulation experts and machinists. By logging transactions in a public blockchain, distributed stakeholders (including regulators) have access to a unified audit trail. However, to outside observers, the anonymized hashes are not interpretable. As AM scales, the expenses of drafting legal paperwork for each license add up, at the disadvantage of small shops who do not retain dedicated counsel. Smart contracts provide a cost-effective andFigure 2.

Smart contracts provide a cost-effective andFigure 2. Schematic of the AMLicense contract, which establishes digital IP rights and licensing terms among distributed stakeholders. End users benefit from independently certifiable data provenance.interpretable alternative.Additive OS Data Management and Directed Acyclic Graphs As discussed in the Introduction, AM presents unique data management challenges. The schema must capture a complex and evolving ontology while enforcing a consistent identification structure. To encourage analytics, the database engine should support sophisticated queries and be accessible through popular programming languages like Python, C++, and MATLAB.

To encourage analytics, the database engine should support sophisticated queries and be accessible through popular programming languages like Python, C++, and MATLAB. Finally, the data importation process should be as automated as possible to minimize human error and ensure usability.5 An exception will be raised if the number of prints does exceeds the allotted amount or the transaction is signed by someone other than the licensee or licensor.259As opposed to relational (SQL) databases, which are relatively inflexible, or RDF-based platforms, which have a steep learning curve, Additive OS is driven by MongoDB, a popular open- source framework for NoSQL (i.e. non-relational) databases [38]. Other than the requirement that each entry has a unique identifier, MongoDB imposes little structure on how documents are stored. Complex objects can be inserted and fetched as human-readable JSON files or language-native representations, like Python dictionaries.

Complex objects can be inserted and fetched as human-readable JSON files or language-native representations, like Python dictionaries. While the Additive OS API is written in Python, data scientists can execute queries in their language of choice.Ontology is captured through an interpretable naming convention that allows data to be encoded as directed acyclic graphs (DAGs). Graphs express adjacency between vertices through edges, and directed graphs specify that edges have a direction from parent to child. Thus, directed acyclic graphs are a certain type of directed graph that does not allow for closed inheritance loops [51]. DAGs are not only an intuitive way to represent AM data, but they introduce certain properties that are useful for mathematical and statistical models, like topological ordering, which can be used to find the shortest path between nodes [52].

In clinical studies, DAGs have been used to help identify biases, confounding factors, and causal relationships between risk factors and diseases [53]. Analogous challenges exist in AM, where process-structure-property relationships can be confounded by complex physics.In Additive OS, the root node of an AM data DAG is a build, which is represented by a unique identifier (UID), e.g. CMU01. This node may include build-specific attributes such as the preheat temperature or links to layerwise images. Leaf nodes that directly inherit the characteristics of the build root (for instance, an individual part) have a UID that includes the build UID, followed by a period delimiter, followed by a secondary index, e.g. CMU01.01. Part objects store parameters such as mechanical testing results or a link to the CAD file.

CMU01.01. Part objects store parameters such as mechanical testing results or a link to the CAD file. Similarly, to encode direction between, for instance, a part and its infill parameters, the infill parameter object for part CMU01.01 would be CMU01.01.IF, and could contain attributes like laser power or hatch spacing. Build ontologies of arbitrary complexity (including heat treatments, machining, and other post- processing steps) can be represented in similar fashion (Figure 3).Figure 3. Data ontology in Additive OS. Build data is imported through flexible spreadsheet templates, stored in MongoDB as JSON documents, and retrieved/visualized as DAGs.260Ideally, build data could be directly imported from build preparation files. However, closed-source machines limit the amount of useful information that can be extracted from these documents.

However, closed-source machines limit the amount of useful information that can be extracted from these documents. Thus, to ensure broad support, Additive OS automatically imports data into MongoDB from flexible, spreadsheet-based templates which were developed as part of a multi-university research project. Once a build tree is imported to a project, process monitoring files, characterization data, and post-processing steps can be attached to the UID that they are associated with. For example, micrographs or yield strength measurements can attached to individual parts, while passive acoustic data may be attributed to the entire build. Additive OS includes an interactive tool for visualizing the build DAGs, which are generated using the PyVis library [54], and they can also be exported in a format compatible with popular graph neural network packages such as PyTorch and Keras.

Under the Data Explorer tab, users can quickly test MongoDB queries on the database before writing code (Figure 4).Figure 4. In addition to the Python API, Additive OS provides tools for inspecting AM data and its structure. The Data Explorer (left) provides a testbed for MongoDB queries before writing code. The DAG visualizer (right) is an interactive and tangible representation of build structure.Peer-to-Peer File StorageGiven the collaborative nature of the AM digital thread, files must have high availability and multiple replications, but security remains paramount. In the cloud computing era, there are many options for reliable, centralized file storage. User-friendly platforms such as Box [55] and Google Drive [56] are popular and intuitive, but difficult to integrate into applications programmatically. Amazon Web Services’ Simple Storage Service (AWS S3) buckets [57] expose this functionality, but coverage is regional.

Amazon Web Services’ Simple Storage Service (AWS S3) buckets [57] expose this functionality, but coverage is regional. If files are not replicated in geographically diverse locations via expensive content distribution networks (CDNs), users can experience high latency when downloading content from monolithic data centers. Most importantly, the aforementioned platforms use name-based addressing schemes, meaning a file is retrieved based on its hierarchically-structured location, e.g. /images/01.png. While this is a human-readable format, it is imprecise. Files can be updated and saved to the same path, either intentionally or maliciously, leading to unnecessary confusion and security risks.The InterPlanetary File System (IPFS) is a decentralized, peer-to-peer CDN that resolves many of the issues with centralized cloud storage platforms [58].

Rather than retrieving content261from remote data centers, clients request file objects from nodes (computers running the client software) that choose to host them, giving preference to those with the lowest latency. Popular files may be locally cached by multiple nodes in different regions, resulting in a scalable, resilient data storage platform capable of hosting photo albums, large scientific datasets, and a copy of Wikipedia [59]. Furthermore, IPFS employs a content-based addressing scheme, where each new object is given a unique identifier that is related to the hash of the content it represents. This serves as a form of version control à la Git [60]. Every file that is added to IPFS through Additive OS undergoes symmetric, client-side encryption using keys managed by AWS [61]. While this is adequate for many situations (and takes full advantage of the scalability of the public IPFS network), it is also possible to create private IPFS networks for additional privacy.

While this is adequate for many situations (and takes full advantage of the scalability of the public IPFS network), it is also possible to create private IPFS networks for additional privacy. As discussed in the next subsection, using existing Additive OS functions, content can easily be replicated among a cluster of dedicated servers for improved availability.Full Customization with Minimal Hardware RequirementsIn contrast to many proprietary platforms, developers can fully customize the Additive OS ecosystem through the open-source API (Figure 5). While the browser-based application exposes the API’s methods through a graphical user interface, the lightweight client can run on low- resource hardware such as ARM devices. For example, inexpensive Raspberry Pi computers can upload process monitoring data directly to the platform using a simple Python script.

For example, inexpensive Raspberry Pi computers can upload process monitoring data directly to the platform using a simple Python script. Data scientists can run sophisticated queries of the MongoDB database to produce curated training sets for machine learning models. Files can be replicated among multiple remote servers to ensure access when primary nodes are offline. We hope Additive OS becomes a collaborative project where contributors continue to improve the platform and add functionality for the benefit of the AM community at large.Figure 5. In addition to the browser-based application, users can use the Additive OS Python library to network their AM digital ecosystem, from data acquisition and organization to file storage and analytics.262Results: Example Use CasesIn this section, two example use cases are presented to show how smart contract functionsare implemented in Additive OS.Use Case #1: SabotageThe first use case is inspired by the threat described in [1].

In this scenario, after a thorough design review, a rotor blade design is approved for manufacture. However, while in transit from the designer to the printing technician, the CAD file is intercepted, and a small defect is embedded in the part. Background studies have shown that even small voids, while difficult to detect upon manual inspection, can be detrimental to part properties [4], [5]. The filenames and sizes of the authentic and sabotaged parts are identical, as only a few bytes were changed. In Additive OS, creating a new project executes the addProject function in the AMProject smart contract (Figure 6), which initiates a project object and assigns a unique project ID. The designer can share the content with the technician by adding their email address to the access list on the project page. When the designer browses for and submits the approved CAD file, the application computes the MD5 checksum, encrypts, and pins the file to the local IPFS node6.

When the designer browses for and submits the approved CAD file, the application computes the MD5 checksum, encrypts, and pins the file to the local IPFS node6. It calls the smart contract’s addHash function, which associates the checksum with this project. Relevant metadata is inserted in the MongoDB database, including the IPFS CID pointing to the file and a URL to view the smart contract transaction on Etherscan [62], which is an independent block explorer for the Ethereum network7. The technician can then download the CAD file from IPFS, decrypt it locally, and quickly compute the checksum using the Validate Data tab on Additive OS or using their own MD5 solver. Even if the MongoDB cluster is compromised or accidentally deleted, the hash of the authentic file will always persist in the smart contract, which has no methods for removing or updating project or hash data.

In fact, all transactions8 will be recorded with the smart contract, ensuring fault-tolerant data provenance.Figure 6. Screenshots of the Additive OS browser interface showing how the application can be used for project organization and tamper detection (use case #1).6 Additive OS does have support for remote pinning through an external service called Pinata [63], but this negates many of the main benefits of operating a node independently. 7 Of course, anyone also has the option to run a their own Ethereum node and view the blockchain data directly for themselves. 8 Read-only functions like getHash are not considered transactions, as they do not alter the Ethereum world state.263Use Case #2: Licensing and IP TheftThe second use case addresses licensing and counterfeiting issues that are commonly discussed by AM practitioners as outlined in [7].

As opposed to the first scenario, this case assumes an adversarial role between a part’s designer and the service bureau that was subcontracted to fabricate it. The designer would like to ensure that the printing shop can only create 5 copies of an unpatented widget, which will be sent to an end customer. In Additive OS, the designer uses a form to create a license object by calling the addLicense function in the AMLicense smart contract (Figure 7). The object stores the ETH wallet addresses of the licensee and licensor, the maximum number of licensed prints, and the checksum of the CAD file. Each time the service bureau prints a part, they use another form to log the operator ID and a hash of the build report. This addPrint transaction creates a tamperproof digital certificate of authenticity, which Additive OS generates a printable QR code for.

This addPrint transaction creates a tamperproof digital certificate of authenticity, which Additive OS generates a printable QR code for. The self-governing smart contract cannot create more licenses than the designer originally specified, so unauthorized copies can be easily detected. Moreover, if the service bureau stole the design for themselves or a competitor, the contract would provide attestation of IP ownership, as the designer could produce the file associated with the checksum supplied in the addLicense transaction.Figure 7. An example workflow for secure print licensing. Each stage invokes a transaction in the AMLicense smart contract to read or write data.264ConclusionsAdditive OS fills a need for an intuitive, open-source AM data and file management platform. Through Ethereum smart contracts built into the backend of the software, our solution integrates sabotage detection and IP theft prevention directly into the user interface.

Through Ethereum smart contracts built into the backend of the software, our solution integrates sabotage detection and IP theft prevention directly into the user interface. For most users, the browser-based application is sufficient for uploading, downloading, sharing, and organizing project files. However, developers and data scientists can enjoy open access to the underlying Python-based API. Data ontology is preserved through a unique ID-based naming convention, which allows users to import build parameters automatically from spreadsheet templates. A flexible MongoDB instance captures the hierarchy of complex DAG structures and provides support for queries in dozens of programming languages. Given the minimal dependencies, the Additive OS client can be run on low-resource hardware, allowing laboratories to connect in situ monitoring equipment or server clusters directly to the network with a few lines of code.

Encrypted peer-to-peer file storage through IPFS enables low-latency, highly available content delivery without compromising privacy.Like many software projects, Additive OS is an ongoing effort that will continuously improve through new features, security updates, and efficiency enhancements. The author is committed to maintaining the repository, updating documentation, and handling issues, but other developers are wholeheartedly encouraged to contribute to this open-source initiative. To try Additive OS in your own facility, you will need a MongoDB cluster (available with a free trial), an AWS account for encryption key management, an Ethereum wallet to cover transaction fees9, and some technical expertise. Documentation is provided on the project’s Github page.References[1] S. Belikovetsky, M. Yampolskiy, J. Toh, J. Gatlin, and Y.

11.L. D. Sturm, C. B. Williams, J. A. Camelio, J. White, and R. Parker, “CYBER-PHYSICALB. Ranabhat, J. Clements, J. Gatlin, K.-T. Hsiao, and M. Yampolskiy, “Optimal SabotageS. B. Moore, W. B. Glisson, and M. Yampolskiy, Implications of Malicious 3D Printer9 The smart contracts used in the examples were deployed to the Ropsten Testnet, which is free to use and suitable for development, but is significantly less stable than the Ethereum Mainnet.265S. Kurpjuweit, C. G. Schmidt, M. Klöckner, and S. M. Wagner, “Blockchain in Additive [7] Manufacturing and its Impact on Supply Chains,” in Journal of Business Logistics, 2021, vol. 42, no. 1, pp. 46–70. [8] [9] [Accessed: 25-Jun-2021]. [10] C. Mandolla, A. M. Petruzzelli, G. Percoco, and A. Urbinati, “Building a digital twin for additive manufacturing through the exploitation of blockchain: A case analysis of the aircraft industry,” Comput. Ind., vol. 109, pp. 134–152, Aug. 2019. [11] N. Papakostas, A. Newell, and V.

ditive/Subtractive Rapid Pattern Manufacturing for Casting Patterns and Injection Mold ToolingMatthew C. Frank, PhD, Frank E. Peters, PhD, Rajesh Karthikeyan Department of Industrial and Manufacturing Systems Engineering Iowa State University, Ames, IA, USAAbstractThis paper presents a Rapid Pattern Manufacturing system that involves both additive and subtractive techniques whereby slabs are sequentially bonded and milled using layered toolpaths. Patterns are grown in a bottom-up fashion, both eliminating the need for multi-axis operations and allowing small features in deep cavities. Similar approaches exist in the literature; however, this system is able to provide a larger range of both materials and sizes, from smaller conventional injection mold tooling to very large wood or urethane sand casting patterns. This method introduces a novel sacrificial support structure approach by integrating a flask into the pattern build process.

This method introduces a novel sacrificial support structure approach by integrating a flask into the pattern build process. The system has been implemented in an automated machine capable of producing patterns in excess of several thousand pounds in a build envelope over 1m3. In this current research, a new layer bonding method using friction stir welding of aluminum plates is presented. In this manner, one can create seam-free laminated aluminum injection mold tooling using a unique combination of industrial adhesives and friction stir spot welding to secure the slab initially, then continuous friction stir welding of layer perimeters that are subsequently machined in a layer-wise process.IntroductionAlthough most rapid prototyping systems are appropriate for testing form, fit and function, they usually require a long processing time; which is reasonable if only one or a few parts are required.

When there is a need to make tens, hundreds or thousands of parts, RP systems are not always the best choice because of the cost and processing time for each part. The availability of rapid prototyping systems for the areas of mass production is limited, but is just starting to see some successes. One of the most commonly chosen manufacturing methods for the mass production of plastic parts is the injection molding process. A wide range of products that vary in their size, shape and complexity can be easily manufactured using injection molding. However, the process of manufacturing an injection mold tool is a complex and highly skilled task that is very costly. Once the design is confirmed it usually takes several weeks or months to actually manufacture and market the product. This is mainly due to the complexity involved in creating the mold tooling.

Once the design is confirmed it usually takes several weeks or months to actually manufacture and market the product. This is mainly due to the complexity involved in creating the mold tooling. There is a strong motivation to implement rapid manufacturing technology for the manufacture of plastic injection molds to reduce the product development time and reduce the cost of manufacturing.Related WorkDifferent types of rapid prototyping and manufacturing methodologies have been developed in the past few decades. Some of the noteworthy methods are Stereolithography (SLA), Fused Deposition Modeling (FDM), Laminated Object Manufacturing (LOM), Selective Laser Sintering (SLS), Electron Beam Melting and other direct metal method, and 3-D Printing. These RP systems are excellent for testing the form, fit and sometimes function of a new design; however most are limited in terms of part accuracy, size and choice of materials.

These RP systems are excellent for testing the form, fit and sometimes function of a new design; however most are limited in terms of part accuracy, size and choice of materials. Hybrid RP process combines the advantage of conventional CNC machining process and a layered manufacturing process to find the solution to these problems [3].Shape Deposition Manufacturing (SDM) was a hybrid process developed at Carnegie Mellon University that employed an additive process to deposit the part or support material using micro-casting process. The material is then machined to achieve the desired accuracy and finish [4]. Solvent welding freeform fabrication technique (SWIFT) creates short run tooling based on solvent welding and CNC machining [5]. For each layer a thin film of high-density polyethylene (HDPE) is printed through a laser printer. HDPE is the solvent mask that prevents unwanted bonding wherever it is applied.

For each layer a thin film of high-density polyethylene (HDPE) is printed through a laser printer. HDPE is the solvent mask that prevents unwanted bonding wherever it is applied. After masking, acetone solvent is applied to the bottom side of the sheet and then stacked to the previous layers and bonded under force. A three axis CNC machine is used to mill down the current sheet to the shape. Computer-aided manufacture (CAM) of laminated engineering materials (LEMs) is another hybrid RP process for fabricating laminated engineering components directly from sheet metal. A laser was used to cut the part slices from the stock of materials such as metals and ceramics. These slices are then assembled together using a selective area gripper [6].

A laser was used to cut the part slices from the stock of materials such as metals and ceramics. These slices are then assembled together using a selective area gripper [6]. However the part accuracy of this system was low due to unpredictable shrinkage which can be as high as 12 – 18 percent [7].Rapid Tooling is an extension of rapid prototyping which is used to prototype mold tooling that can be used for early production. Rapid tooling (RT), allows manufacturing of production tools such as molds and dies rather than the final part itself which can reduce the lead time for the product to reach the market [2]. Rapid laminated tooling is similar to laminated object manufacturing (LOM), In the LOM process, each layer of the part is formed from an adhesive coated sheet of paper which were subsequently cut with a laser [10]. Instead of paper, other forms of laminated tooling used sheets of metals. These sheets of metals could be joined together by bolts, welding or brazing.

Instead of paper, other forms of laminated tooling used sheets of metals. These sheets of metals could be joined together by bolts, welding or brazing. Extensive research has been conducted on creating tooling for plastic and metal forming processes. Laminated tooling is not a new concept, where research and development in this field has been conducted since early researchers like Nakagawa back in 1980, who were creating blanking dies for sheet metal components by using bainite steel sheets for the tool face and cheaper steel as backing plates. The steel sheets were cut using a laser, stacked horizontally and joined together by using mechanical fasteners [11].Most of these laminated tool manufacturing processes follow a build sequence of cut, stack and bond. First, the plates are cut to the required cross section using a laser or EDM process and then these laminates are cleaned and stacked in either horizontal or vertical orientation. Finally, the stacked plates are bonded together.

Finally, the stacked plates are bonded together. Many researchers used different bonding methods, such as mechanical fasteners, laser welding, diffusion bonding and bonding by adhesives. The more popular joining method being the use of mechanical fasteners such as bolts and rivets to join the laminates together [11-14]. However, most of these processes do not provide a complete automated process planning solution. In addition, selecting the thickness of the laminates has always been an issue, where selecting thin laminate thickness of 0.5 and 2mm increased both the complexity and time in creating the toolingThe proposed process, Rapid Manufacturing of Plastic Injection Mold (RMPIM) uses a build sequence of stacking-bonding-cutting of aluminum plates as opposed to cutting-stacking-bonding cited in most of the literature. This approach should more readily enable completely automated process planning for creating injection mold tooling.

This approach should more readily enable completely automated process planning for creating injection mold tooling. The proposed process uses a new layer bonding method; a unique combination of industrial adhesives and friction stir welding. A hybrid Rapid Pattern Manufacturing system (RPM), previously developed by the authors, can create large wooden casting patterns [8-9]. The process combines depositing a thick slab of Medium-density fiberboard (MDF) and a three axis CNC machine to cut the board to a defined layer thickness and to create part geometry on the layer. The advantage of this system is that the patterns are built in the bottom-up fashion so a small tool can be used to mill deep cavities without the use of multi-axis (beyond three-axis) CNC machines.Friction stir welding (FSW) is a solid-state joining process.

A non-consumable rotating tool with a specially designed pin and shoulder is inserted into the abutting edges of sheets or plates to be joined and traversed along the line of joint. Frictional heating is produced from rubbing of the rotating shoulderwith the work pieces, while the rotating pin causes plastic deformation of work piece. The heating is accomplished by friction between the tool and the work piece and plastic deformation of work piece. The localized heating softens the material around the pin and the combination of tool rotation and translation leads to movement of material from the front of the pin to the back of the pin where it is forged into a joint [15-16].(a)(b)Figure 1 - (a) Schematic diagram of friction stir welding [16] (b) FSW process [17]Overview of ProcessThe proposed rapid mold manufacturing system uses a hybrid manufacturing method, a combination of additive and subtractive process to create plastic injection molds.

The basic process involves adding a layer of aluminum plate, which is then subsequently machined to obtain the 3D shape of that particular layer. This process uses friction stir welding for layer bonding, which would enable the creation of seam-free laminated aluminum injection mold tooling. The fundamental additive and subtractive nature of the process is illustrated in figure 2 for simply two layers.When a new plate is added to the base plate, it needs to be clamped for the subsequent welding and machining process, but the use of mechanical fixtures and clamps for machining in rapid will manufacturing potential create problem for collision of the tool/spindle and the work piece setup.

Therefore, the process proposed in this paper uses a combination of industrial adhesives and friction stir spot welding to secure the aluminum plates.FSW Toola(b) Cross sections are bonded using Friction Stir Welding(A) New layer of plate is deposited(c) CNC machiningThe adhesive is applied in the areas of both the boundaryFigure 2 – Basic process steps using f riction stir welding f or layer bonding (additive) and CNC machining (subtractive) f or 3D layer shapingwall cross section and the mold cross section. The boundary wall in this process acts as a sacrificial support structure which aids in orienting and fixturing of the deposited plate. The adhesively bonded plate is then spot welded using friction a stir spot welding process (FSSW). This is because the strength of the adhesives alone is assumed insufficient withstand the forces of the continuous friction stir welding process (FSW).

This is because the strength of the adhesives alone is assumed insufficient withstand the forces of the continuous friction stir welding process (FSW). Both the adhesives and the spot welding acts as a clamp so that the plates will not move or shear due to friction stir welding and generally keeps the plate flat and undistorted.A face milling operation is performed on the deposited plate prior to friction stir spot welding to make sure that the plate is flat and parallel with the work table of the machine. The plate can then be pre- drilled in all the spot weld locations and at the entry point of the continuous friction stir welding process. The pre drilled holes will reduce the force acting on the mold work piece and the machine table from the friction stir spot welding and friction stir welding processes [19].

The pre drilled holes will reduce the force acting on the mold work piece and the machine table from the friction stir spot welding and friction stir welding processes [19]. Next, the plates are bonded together using a continuous friction stir welding process so that the tooling can withstand the pressure of injection molding process and to create a seam-free tool.Lastly, the plates are machined using ball and flat end mills to obtain the part geometry of that particular layer. The plate is once again face milled to remove burrs from the friction stir welding process so that the next plate can be stacked onto a known height flat surface. The process of stacking, bonding and machining of the plate is continued sequentially until the complete mold tooling defined by the CAD model is created. For previous work in sand casting patterns, the outer boundary wall would either be discarded or kept as the flask if draft was designed into the wall geometry.

For previous work in sand casting patterns, the outer boundary wall would either be discarded or kept as the flask if draft was designed into the wall geometry. For aluminum tooling, the wall would simply be removed at the end. It should be noted that female mold tooling, having a boundary region already, would not necessarily need the boundary wall added to the CAD model.

It should be noted that female mold tooling, having a boundary region already, would not necessarily need the boundary wall added to the CAD model. To illustrate the step by step process more clearly, Figure 3 illustrate a few layers of a hypothetical piece of tooling as the process steps through plate addition, Friction Stir spot and cross section welding, and then CNC machining.Initial SlabAdhesive depositedNew layer appliedFriction Stir Spot WeldingFriction Stir Welding of Cross SectionCNC machining of 3D layer shape…Adhesive for next layer……New slab…Completed buildTooling with boundary wall removedFigure 3 – Detailed process steps for the proposed rapid tooling systemProcess Planning MethodA critical characteristic for any rapid prototyping and manufacturing system is completely automated process planning. That is, the process must be able to be executed from a CAD model, with little or no human intervention or skill required.

That is, the process must be able to be executed from a CAD model, with little or no human intervention or skill required. The main problems that need to be addressed in order for this process to be completely automated are:Determining the boundary wall dimension • Finding the Number, Location and Sequence of Friction Stir Spot Welds (FSSW) • Toolpath planning for Friction Stir Welding (FSW) • Toolpath planning for CNC machiningThe following section presents methods that will enable the automated process planning for this system, using only a CAD model and basic system and processing parameters.Determining the Boundary Wall DimensionThe aluminum plates that are added layer by layer to create the tooling need to be oriented and clamped together for friction stir welding and subsequent machining process.

The use of fixtures and clamps for friction stir welding or machining in rapid manufacturing create a potential problem for collision of the tool/spindle and the work piece setup. Therefore, in this process a combination of industrial adhesives and friction stir spot welding is used to secure the aluminum plates for machining. Each pattern is enclosed by a boundary wall that serves as a sacrificial fixture to aid in the layer based addition of plates. When a new plate is added to the part, the plate can be positioned securely and repeatedly with the support of boundary wall. Initially, an adhesive will hold the plates together and be able withstand the forces from the subsequent FSSW process without the shearing of plates.The intent is to secure the plate with enough adhesive strength to enable spot welding, which then enables continuous friction stir welding of the pattern geometry within the plate.

Obviously, if the load acting on the adhesive during spot welding is more than the strength of the adhesive; the bond fails. Adhesive joint strength can be increased by increasing the area of the bond, (e.g. doubling the bond area approximately doubles the force required for failure stress). This paper presents a method to determine the dimension of the boundary wall (Figure 4a) based on size of the mold geometry, forces acting on the plate due to friction stir spot welding, mechanical properties of the adhesive used and the boundary wall clearance. The boundary wall clearance (dimension a in Figure 4b) is the required space between the boundary wall and mold geometry.

The boundary wall clearance (dimension a in Figure 4b) is the required space between the boundary wall and mold geometry. The objective of adhesive bonding is such that it will transfer and distributes the stress so the plates will not shear due to the forces of the friction stir spot welding process.L2(a)(b)L1wAdhered boundary box (although adhesive elsewhere )Figure 4 – Mold geometry showing lengths, L1,2 and width, W of the boundary wallThe polygon cross section of the CAD model differs from slice to slice but the size of the boundary wall is constant throughout. Therefore, the dimension of the boundary wall is determined such that the area of adhesive applied to the boundary wall should be able to withstand the forces from the the extreme points FSSW process.

The area of the bond to withstand the forces will depend on the mechanical properties of the adhesives and the stress acting on it. The stresses in adhesives arising from the differential shear strain were analyzed by Volkersen [20].

Therefore the plates are subsequently bonded together using a Friction Stir Spot Welding process. The load acting on the spot welds due to FSW process will not be uniform and if the load acting on a particular spot weld is greater than the failure load, (cid:1858)(cid:3033), the spot weld will fail. Therefore, an algorithm will determine the location and number of spot welds needed so that load acting on each of the spot welds is less than the failure load. The algorithm will consider inter layer dependency; when any spot weld location is same as the location of a spot weld or exit hole location of a previous layer, then the spot weld location must change. The location will be moved on the offset curve by a distance 2r, where r is the radius of the FSW tool. This is because the friction stir welding process will leave a hole at the retracting point of the tool; previously mentioned as the exit hole.

This is because the friction stir welding process will leave a hole at the retracting point of the tool; previously mentioned as the exit hole. The shear force (cid:1858)(cid:3046) acting on the plate will cause mode II type failure (sliding mode), in plane overlap shear failure. The failure rule for the spot weld for mode II is given by the equation,(cid:4666)(cid:1858)(cid:3046)/(cid:1832)(cid:3046)(cid:4667)(cid:1503) (cid:3409) 1(10)is an unknown that will define the failure relation between independent modes. For any single loading, regardless of the , the equation will satisfy the failure condition. It means that when the applied load reaches value ofThe denominator (cid:1832)(cid:3046) represents the shear strength of the spot weld. The value ofcan be the strength of the spot weld, the spot weld will fail for each single load [21].

The value ofcan be the strength of the spot weld, the spot weld will fail for each single load [21]. The value of determined by experiments; for example, tests conducted by Wang [22] gave the value of (cid:1503) to be 2 for small thickness to radius ratio (thickness of plates to radius of the weld). In this thesis, the t/r ratio is small so the value of (cid:1503) is taken as 2. In future work, when this process uses thick plates the value of (cid:1503) should be determined by experiments.Friction Stir Spot Welding in the boundary wall cross section:For each layer to be spot welded to the previous layer, the possible regions for the location of the spot are the boundary wall and the cross section area of the polygon of that slice. This is because in the previous layer all the regions of the plate except for the polygon cross section of the slice and the boundary wall will be machined.

This is because in the previous layer all the regions of the plate except for the polygon cross section of the slice and the boundary wall will be machined. There will be one spot weld in each of the boundary wall sides which will act as a clamp to hold the current plate to the previous plates. The boundary wall used in this thesis will only be in rectangular shape (four sided), therefore there will be four spot welds in each layer which will be either in the corners of the boundary wall or in the mid span of the boundary walls. The location of the spot welds is alternated for subsequent layers as in figure this is because of the exit hole in the previous layer.424(a)1(b)2133Figure 5 – Friction stir spot welding (a) on layer(cid:2924) (b) on layer(cid:2924)(cid:2878)(cid:2869)The sequence of spot welds for the rectangular boundary wall is as shown in Figure 5.

In figure 5a, the location of the first spot weld will be on the bottom most point of the left boundary wall side and the second spot is the point diagonal to the first spot weld. The location of the second spot weld is selected such that it increases the moment arm from the first spot weld so the force acting on the spot weld is reduced. The location of the third and fourth spot weld is as shown in the figure. Similarly in figure 5b, the location of the second spot weld is selected such that it has increased moment arm from the first spot weld.

Similarly in figure 5b, the location of the second spot weld is selected such that it has increased moment arm from the first spot weld. This alternating pattern should generally provide a robust and flat outer wall boundary to ensure repeatable placement of each new plate, not affected by build height.Friction Stir Spot Welding in the mold cross section:The four spot welds on the boundary wall are to clamp the plates firmly so that it can withstand the forces from friction stir welding process without the shearing of plates, but the forces involved infriction stir welding are generally of high magnitude therefore spot welds are also needed in the polygon of mold cross section. In addition, it is imperative that the plate is in intimate contact with the layer below before continuous FSW of the layer boundary; therefore spot welds are needed in the mold cross section.

In addition, it is imperative that the plate is in intimate contact with the layer below before continuous FSW of the layer boundary; therefore spot welds are needed in the mold cross section. The four spot welds on the boundary wall cross section will be same throughout the mold but the number of spot welds on the mold cross section polygon will depend on the size of the mold. The location of the spot welds in mold cross section polygon will depend on the critical number of welds, (cid:1866)(cid:3030)(cid:3045)(cid:3036)(cid:3047)(cid:3036)(cid:3030)(cid:3028)(cid:3039). (cid:1866)(cid:3030)(cid:3045)(cid:3036)(cid:3047)(cid:3036)(cid:3030)(cid:3028)(cid:3039) for a mold cross section polygon is determined such that it prevents the plate from shearing during friction stir welding and also the load acting on each of the spot weld (cid:1858)(cid:3046) is less than the failure load (cid:1858)(cid:3033).

The location of the sixth and subsequent spot welds will depend on the region analysis of the previous spot welds.

Within that region the midpoint of the polygon section, M will be determined as the location of sixth spot weld. This region-based analysis will continue until the loading condition is satisfied. The minimum distance between two spot welds should be at least 2r, where r is the radius of the smallest friction stir welding tool available in the machine tool library. When a particular mold cross section polygon does not have sufficient space to accommodate all the spot welds required to with stand the forces from FSW process, then FSW cannot be performed in that layer. This problem could occur at the peaks of tall thin structures, but will not be formally addressed in this paper.Tool Path Planning for Friction Stir WeldingThe aluminum plates are oriented, fixtured and clamped using adhesives and friction stir spot welding. The plates are then welded together using continuous friction stir welding process.

The plates are then welded together using continuous friction stir welding process. The FSW process welds the two plates together which is the additive process then it is subsequently machined in a subtractive process for creating the final 3D shape of the mold tooling. The FSW path is generated such that it moves along the perimeter of the layer polygon so it creates seam-free laminated aluminum injection mold tooling. The FSW tool path on each layer will simply depend on the polygon profile on (cid:1871)(cid:1864)(cid:1861)(cid:1855)(cid:1857)(cid:3041)(cid:2878)(cid:2869) as shown in the figure. This is because if the tool path is based on any other polygon profile then the FSW tool will affect the previously machined layer.

The friction stir welding process will leave a hole at the exit point it is called exit hole so the exit point of the FSW cannot be same for all the layers (or at least cannot repeat in the same x-y location on the immediate next layer. Therefore the exit point of the subsequent layers will be offset from the entry points of the previous layers at least by the diameter of the FSW tool. In addition it is advantageous to move the exit hole toward the cross section interior, as shown in Figure 7, thereby burying the void inside the aluminum mold geometry.The last step in process planning is the 3D CNC machining toolpath planning. This step is exceedingly straightforward and will not be presented in this paper.

This step is exceedingly straightforward and will not be presented in this paper. Essentially, each layer is face milled to a flat surface, then executed upon by waterline toolpaths using a flat and ball end mill for roughing and finishing, respectively.Preliminary ResultsExperiments with relatively thin aluminum layer bonding have been conducted in order to evaluate the possibility of using Friction Stir Welding as the bonding method for a rapid tooling system. A tool was created from H13 tool steel, based on the design of a Flared-Triflute developed by TheWall boundary(a)FSSW(b)Exit holeEntry point(c)(d)Figure 8 – Test sample (a) layer sample af ter all steps, (b) close up of cross section geometry and exit hole, (c) FSW tool geometry and (d) example FSW f rom entry to exit holeWelding Institute (TWI), UK. The tool had a pin diameter and length of 6mm and 4.7mm, respectively, and a shoulder diameter of 18mm.

The tool had a pin diameter and length of 6mm and 4.7mm, respectively, and a shoulder diameter of 18mm. Aluminum plates at 1/8” thick were bonded using Friction Stir Spot Welding and then continuous Friction Stir Welding. The plunge and retract federates were 7 and 11inches per minute, while the continuous stirring was conducted at 1800rpm and 25inches per minute. Lastly, the cross sectional geometry of the layer was machined using a face mill and flat-end mill. Images of the tooling, sample layers and a friction stir weld are shown in Figure 8.Conclusion and Future WorkThis work illustrates a new method for bonding aluminum layers for rapid tooling, which could enable high performing rapid tooling based on a hybrid approach. The work is a further step from a previously developed system called RPM, or Rapid Pattern Manufacturing. The RPM system was developed for the rapid construction of mainly wood patterns for steel casting sand-mold making.

The RPM system was developed for the rapid construction of mainly wood patterns for steel casting sand-mold making. RPM automatically generates process plans from a CAD file alone, and then sequentially stacks, bonds and machines layers of wood to create very large patterns. In order to use aluminum as a possible material, enabling tooling such as injection molds, a new method for bonding layers needs to be developed. This work showed preliminary studies using a combination of industrial adhesives, Friction Stir Spot Welding and Friction Stir Welding. The system is intended to use adhesives to initially secure the aluminum plates for spot welding, which in turn, enables continuous friction stir welding of the tooling cross sectional contours. Once bonded, the RPM system using a 3-axis milling apparatus to create accurate 3Dcontoured shapes. The future of this work is envisaged as seam-free, quasi monolithic aluminum tooling.

Once bonded, the RPM system using a 3-axis milling apparatus to create accurate 3Dcontoured shapes. The future of this work is envisaged as seam-free, quasi monolithic aluminum tooling. The concept of quasi-monolithic is based on the idea that the friction stir welding on the cross sectional contours could be executed near the edge of each lamination; hence, the subsequent CNC machining would actually mill through this welded boundary. Viewed from above, the stack of aluminum plates created by the RPM system would appear to be one continuous aluminum “shell” surface (Figure 9). Within the tooling, one could bury all exit and entry holes from welding, the interlaminate spaces where adhesive remains, and perhaps even integrated cooling channels. Moreover, this approach would allow for extremely deep cavity machining of complex geometry with no collision conditions.

Moreover, this approach would allow for extremely deep cavity machining of complex geometry with no collision conditions. Not only could tooling be created in a rapid fashion, but this process could enable revolutionary capabilities and performance, along with the accuracy of a CNC machined surface. The efficacy of this approach will require extensive continued testing and research to evaluate weld capabilities, strength, geometric limitations, etc.(a)(b)(c)Figure 9 – Proposed Seam Free Laminated tooling (a) two layers Friction Stir Welded, (b) layers af ter machining through prof ile welds, and (c) illustration of a seam-f ree tooling stack up, with laminations, exit holes, cooling channels, etc contained within the tool surf aceAcknowledgementResearch was sponsored by the U.S. Army Benet Laboratories and was accomplished under Cooperative Agreement Number W15QKN-06-R-0501.

Army Benet Laboratories and was accomplished under Cooperative Agreement Number W15QKN-06-R-0501. The views and conclusions contained in this document are those of the authors and should not be interpreted as representing the official policies, either expressed or implied, of U.S. Army Benet Laboratories or the U.S. Government. The U.S. Government is authorized to reproduce and distribute reprints for Government purposes notwithstanding any copyright notation heron.Reference[1] Rafiq Noorani, “Rapid prototyping : principles and applications”, 2006 John Wiley & Sons, Inc[2] N.P. Karapatis, J.-P.S. van Griethuysen and R. Glardon, “Direct rapid tooling: a review of current research”, Rapid Prototyping Journal, Volume 4 · Number 2 · 1998 · pp.

DebRoy and H. K. D. H. Bhadeshia, “Recent Advances in Friction Stir Welding – Process, Weldment Structure and Properties”, Progress in Materials Science 53 (2008) 980–1023[18] M. Song, R. Kovacevic, “Thermal modeling of friction stir welding in a moving coordinate system and its validation”, International Journal of Machine Tools & Manufacture 43 (2003) 605–615[19] Terry Khaled, “An ouysider looks at Friction Stir Welding”, Report#: ANM-112N-05-06, July 2005[20] Adams, Robert D, “Structural adhesive joints in engineering”, London; New York: Chapman & Hall, 1997[21] P.Wung, “A force-based failure criterion for spot weld design”, Experimental Mechanics, Volume 41, Number 1 / March, 2001[22] Case, John, The strength of materials; a treatise on the theory of stress calculations for engineers, London, E. Arnold & co., 1925.

Microstructures before and after HTs are analyzed using both optical and scanning electron microscopy (SEM). Both quasi-static uniaxial tensile and hardness tests are performed to measure mechanical properties. The tensile results indicate that the non-heat treated (NHT) condition possesses higher strengths but lower ductility as compared to the other HT conditions (i.e., SR, SA, HIP, SR+SA, and SR+HIP). By employing the two-step HT conditions (i.e., SR+SA and SR+HIP), no significant changes on tensile properties as compared to the individual single-step HT conditions (i.e., SA or HIP) are observed.

By employing the two-step HT conditions (i.e., SR+SA and SR+HIP), no significant changes on tensile properties as compared to the individual single-step HT conditions (i.e., SA or HIP) are observed. The findings suggest that two-step HTs are not required for LP-DED 316L SS unless HIP is needed to minimize volumetric defect contents.Keywords: 316L stainless steel; Laser powder directed energy deposition; Heat treatment; Microstructure; Tensile propertiesIntroductionLP-DED is an additive manufacturing (AM) method that uses injected metal powder flow as feedstocks and the laser beam as an energy source to deposit the powder materials layer by layer to obtain the near net-shaped components [1].

The LP-DED technique has some benefits compared to other AM methods such as fabrication of relatively larger parts, repair capabilities of damaged or faulty components, the ability to clad different materials on existing parts , and reduced material waste [1][2].Austenite SS such as 304L SS and 316L SS is commonly used SS in numerous industrial applications due to their great corrosion resistance, oxidation resistance, and formability [3,4]. Recently, they have received a lot of attention from AM community. Many researchers have investigated the effects of processing parameters including laser power, scan speed, and hatching distance on microstructure and mechanical properties [5,6]. Even with optimized process parameters, the AM 316L SS has shown anisotropic mechanical properties and corrosion resistance which could be attributed to the characteristic thermal history experienced during AM397fabrication processes which accordingly could affect the alloy’s microstructure [7].

One common way used in the community to reduce the anisotropic properties in AM components is post-process HT [8][9]. However, due to the AM’s unique thermal history, the standard HTs may or may not achieve similar effects on microstructures and mechanical properties of the AM alloys as they do on conventionally processed counterparts.Two-step or multi-step HTs become common practices for AM components to resolve anisotropic properties due to mainly rapid cooling in AM process. The SR is considered a first step for the AM parts to remove the residual stresses induced during fabrication processes. SR is usually conducted before removing the parts from the build plate. After SR, conventionally standardized HTs such as annealing and aging for wrought materials are also utilized in AM materials to homogenize and enhance microstructures and mechanical properties, respectively.

Furthermore, the HIP process is frequently used for AM materials since it has been proved to be beneficial effect on closing up certain types of defects induced during AM fabrication processes [10].This study examines the effect of such two-step HTs on 316L SS fabricated by LP-DED technology. Three different two-step HTs such as SR at 899°C + standard SA, SR at 1010°C + standard SA, and SR at 899°C + standard HIP were considered. All three two-step HTs were further compared with single-step HTs (i.e., SR at 899°C, standard SA, or standard HIP). The effects of HTs were studied on microstructure as well as on uniaxial tensile behaviors of the alloy. This article is presented in the following sequences. First, the experimental procedures such as materials preparation and testing are explained.

This article is presented in the following sequences. First, the experimental procedures such as materials preparation and testing are explained. Then, the experimental results are presented and discussed; and finally, conclusions are drawn.Experimental ProcedureThe gas atomized virgin 316L SS powders used to fabricate specimens were provided by the Carpenter Additive. Table 1 lists the chemical composition of the 316L SS powders along with the nominal composition of the alloy according to ASTM A240/UNS standard [11]. LP-DED 316L SS cylindrical rods (15.24 mm Dia.

Table 1 lists the chemical composition of the 316L SS powders along with the nominal composition of the alloy according to ASTM A240/UNS standard [11]. LP-DED 316L SS cylindrical rods (15.24 mm Dia. × 104.14 mm height) were fabricated vertically (i.e., parallel to loading direction) using the following parameters: laser power of 1070 W, layer height of 381 µm, travel speed of 1016 mm/min, and powder feed rate of 15.20 g/min.Table 1 The chemical composition of 316L SS powder used in this study along with the nominal composition of the alloy according to ASTM standard A240/UNS [11].Nominal (ASTM A240/UNS)MeasuredMinMaxFe Cr Ni Mo Mn Si N16.0 10.0 2.00Balance18.0 14.0 3.00 < 2.00 < 0.75 < 0.1065.80 17.73 12.70 2.41 0.58 0.62 0.10398C P S< 0.030 < 0.045 < 0.0300.02 0.007 0.006The cylindrical rods were heat treated using various HT schedules as listed in Table 2.

Different variables used for the HTs in this study such as temperatures and soaking times of SR and HIP were designed according to the standard AMS 2759 [12] and ASTM F3301 [10] procedures. In addition, extra coupons (15.24 mm Dia.

In addition, extra coupons (15.24 mm Dia. × 3 mm thickness) were heat treated separately for further investigation to examine the effects of holding temperatures ranging from 600°C to 1200°C on microstructures.Table 2 Details of different heat treatment recipes.Heat treatments NHTHT01SR899°C / 2 hours / Air coolingSAHIPHT02HT03HT04899°C / 2 hours / Air cooling 1010°C / 2 hours / Air cooling1100°C / 2 hours / Argon quenching 1100°C / 2 hours / Argon quenching 1100°C / 2 hours / Argon quenchingHT05HT06899°C / 2 hours / Air cooling1163°C / 3 hours / 103 MPa / Argon quenching 1163°C / 3 hours / 103 MPa / Argon quenchingAccording to ASTM E3 [13], small coupons from specimens listed in Table 2 and extra coupons were prepared for the standard metallographic characterization.

Microstructures of all HTs were analyzed by KEYENCE VHX-6000X optical microscope and Zeiss Crossbeam 550 scanning electron microscope (SEM) equipped with electron backscatter diffraction (EBSD) and elemental diffraction spectroscopy (EDS) detectors. The microstructural analysis was performed on both normal and transverse planes (i.e., perpendicular and parallel to the build orientation, respectively) for all HT coupons. For optical microscopy, heat treated small coupons were chemically etched using HCL:HNO3:CH3COOH etchant to reveal microstructure and meltpool [14]. EBSD scans were conducted using followed configurations: magnification of 50X, voltage of 15 kV, current of 200 pA, and step size of 4µm. The average grain sizes and annealing twins were measured from the EBSD data.

The average grain sizes and annealing twins were measured from the EBSD data. It should be noted that the average grain sizes were calculated excluding the ∑3 twin boundaries to avoid miscalculation of twins as grains.Upon HTs, cylindrical rods were machined to the final geometry of tensile test specimens specified the ASTM E8 standard [15], as shown in Figure 1. Uniaxial tensile tests were conducted on test specimens with HTs listed in the Table 2 using a strain rate of 0.001 mm/mm/s according399to the ASTM E8 standard procedure [15]. An extensometer was used during the initial stage of tensile tests until a 0.045 mm/mm strain for accurate determination of yield strength. Upon removal of the extensometer, the tests were resumed using the displacement-controlled mode.

Upon removal of the extensometer, the tests were resumed using the displacement-controlled mode. The average values reported for the tensile properties, i.e., the average ultimate tensile strength (UTS), yield strength (YS), and elongation to failure (EL), were calculated over two test data.Figure 1 Geometry of tensile test specimens according to ASTM E8 [15].Results and DiscussionThe microstructures of NHT and HT coupons were analyzed by optical microscopy and are presented in Figure 2 and Figure 3, respectively. The melt pool boundaries and fine dendritic microstructures, which are the characteristics of the AM materials [14], are visible in the normal and transverse planes of NHT specimens. As shown, the melt pool boundaries and dendritic microstructures have not been fully removed until 900 °C. The coupons heat treated by temperatures above 1000 °C do not present melt pool boundaries and fine dendritic microstructures anymore.

The coupons heat treated by temperatures above 1000 °C do not present melt pool boundaries and fine dendritic microstructures anymore. It indicates that some recrystallization and grain growth has occurred.The inverse pole figures (IPF) and grain boundaries obtained from EBSD analysis are shown in Figure 4 for all HTs. As shown, increasing the temperature of the first step of HTs up to 1000°C does not show significant effects on average grain sizes. In addition, the number of annealing twins that are calculated using the percentage of ∑3 coincidence site lattice (CSL) boundaries is not changed by HTs using temperatures below 1000°C. Similar observations were reported from wrought 316L SS counterparts in the literature [16]. The average grain size and annealing twin fraction are increased when the HT temperatures are higher than 1000°C.

The average grain size and annealing twin fraction are increased when the HT temperatures are higher than 1000°C. It again suggests that recrystallization and grain growth have been caused.400Figure 2 Optical micrographs of LP-DED 316L SS in NHT condition obtained on normal plane in (a) and transverse plane in (b) (i.e., parallel and perpendicular to the build orientation, respectively).401Figure 3 Optical micrographs of LP-DED 316L SS heat treated by holding temperatures of (a) 600°C, (b) 700°C, (c) 800°C, (d) 900°C, (e) 1000°C, (f) 1100°C, and (g) 1200°C for 2 hours.402Figure 4 Inverse pole figure (IPF) and grain boundaries of specimens after heat treated by different temperatures: (a) NHT, (b) 600°C, (c) 700°C, (d) 800°C, (e) 900°C, (f) 1000°C, (g) 1100°C, and (h) 1200°C.The tensile results, including the stress-displacement curves, and bar graphs of all specimens are presented in Figure 5.

Furthermore, the details of tensile properties such as the ultimate tensile strength (UTS), yield strength (YS), and elongation to failure (EL) of each condition are listed Table 3. As shown in Figure 5 and Table 3, the NHT specimen has shown the highest UTS of 581.09 MPa and YS of 398.65 MPa and the relatively low EL of 81.25%. Tensile properties LP-DED 316L SS fall the mechanical test requirements proposed by ASTM A240 standard of the wrought 316L SS counterparts that are presented as dotted lines in Figure4035(b) [11]. However, UTS and YS are lower than the laser beam powder bed fusion (L-PBF) 316L SS (UTS of 612 MPa and YS of 490 MPa) regardless of HT conditions [8].Figure 5 Uniaxial tensile behavior of LP-DED 316L SS specimens in NHT and different HT conditions: (a) stress-displacement curves and (b) bar chart of UTS, YS, and EL for each HT condition.

ASTM A240 mechanical test requirements are presented as dotted lines in (b).Table 3 UTS, YS, and EL of specimens after each heat treatment.Heat treatment NHT (AB) HT01 (SR) HT02 (SA) HT03 (SR+SA) HT04 (SR+SA) HT05 (HIP) HT06 (SR+HIP)UTS (MPa) 581.09 578.36 554.91 553.85 551.81 554.18 557.48YS (MPa) 398.65 314.56 272.43 274.53 272.76 249.86 254.52Elongation (%) 81.25 78.75 81.25 90.00 87.50 86.25 91.25As shown earlier, the SR at 899°C for 2h (HT01) does not have significant effects on the microstructure of LP-DED 316L SS specimens as compared with NHT conditions. It could be404attributed to the low temperature, which subsequently resulted in similar UTS and YS in SR (HT01) as compared to the NHT specimen. Except for the HT01 specimens, all the other HTs have shown similar UTS and YS, which were lower than those of the NHT specimen.

Except for the HT01 specimens, all the other HTs have shown similar UTS and YS, which were lower than those of the NHT specimen. As discussed earlier in the previous section, this could be ascribed to the effect of HT temperature during the first step on microstructure, resulting in better homogenization and removal of microsegregation and dendritic microstructures which may have some strengthening effect [17].As shown in Figure 5, the two-step HT conditions such as SR at 899°C/2h + SA (HT03), SR at 1010°C/2h + SA (HT04), and SR at 899°C/2h + HIP (HT06) do not show significant changes on tensile properties as compared to the single-step HT conditions (i.e., SA (HT02) and HIP (HT05)). It implies that standard SA without SR can bring reasonable microstructural homogenizations resulting in insignificant effects on tensile properties. However, HIP is commonly required to reduce volumetric defects of AM material [10].

However, HIP is commonly required to reduce volumetric defects of AM material [10]. Therefore, HIP is still important for fatigue critical applications because the volumetric defects can be the major cause of the early fatigue crack initiations resulting in short fatigue lives [18].ConclusionsIn this study, the effects of different heat treatment conditions on microstructure and uniaxial tensile properties of LP-DED 316L SS were investigated.

The following conclusions could be drawn on the experimental observations:The melt pool boundaries and fine-scale cellular structures were not fully removed when the holding temperature was below 1000 °C, due to the sluggish diffusion.Upon HT at higher temperatures than 1000°C, the average grain size and annealing twin fraction increased, suggesting that some recrystallization/grain growth occurred.Tensile properties of LP-DED 316L SS in non-heat treated consistent with those of mechanical test requirements dictated by the ASTM A240 standard; they showed slight inferiority as compared to the L-PBF 316L SS.Tensile strengths of LP-DED 316L SS in NHT were the highest among all HT conditions which were traded off with the relatively low elongation to failure.AcknowledgementThis research is partially supported by the National Aeronautics and Space Administration NASA under Cooperative Agreement No. 80MSFC19C0010. This paper describes objective technical results and analysis.

ASTM Stand 2018:3. https://doi.org/10.1520/F3301-18A.2.[11] American Society for Testing and Materials. ASTM A240: Standard Specification for Chromium and Chromium-Nickel Stainless Steel Plate , Sheet , and Strip for Pressure Vessels and for General Applications. ASTM Int 2004;I:12. https://doi.org/10.1520/A0240.[12] Committee AMSEC and LAS. Heat Treatment of Steel Parts General Requirements 2019.https://doi.org/https://doi.org/10.4271/AMS2759G UI - AMS2759G.[13]Interanational A. Standard Guide for Preparation of Metallographic Specimens Standard Guide for Preparation of Metallographic Specimens 1. ASTM Int 2012;03.01:1–12. https://doi.org/10.1520/E0003-11R17.1.[14] Nezhadfar PD, Shamsaei N, Phan N. Enhancing ductility and fatigue strength of additively manufactured metallic materials by preheating the build platform. Fatigue Fract Eng Mater Struct 2021;44:257–70. https://doi.org/10.1111/ffe.13372.[15] ASTM International.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceADDITIVELY MANUFACTURED CONFORMAL NEGATIVE STIFFNESS HONEYCOMBSD. A. Debeau*, C. C. Seepersad*Department of Mechanical Engineering, The University of Texas at Austin, Austin, TX 78712AbstractThis study investigates the static and dynamic mechanical performance of conformalnegative stiffness honeycomb structures. Negative stiffness honeycombs are capable of elastically absorbing a static or dynamic mechanical load at a predefined force threshold and returning to their initial configuration after the load is released. Most negative stiffness honeycombs rely on mechanical loading that is orthogonal to the base of the structure. In this study, a more three dimensional design is presented that allows the honeycomb to conform to complex surfaces and protect against impacts from multiple directions.

In this study, a more three dimensional design is presented that allows the honeycomb to conform to complex surfaces and protect against impacts from multiple directions. The conformal designs are additively manufactured in nylon and stainless steel and subjected to quasi-static mechanical loading and dynamic mechanical impact tests that demonstrate their impact protection capabilities.IntroductionA honeycomb design that absorbs energy and mitigates mechanical impacts elastically is very desirable because it can recover from an initial impact and respond to a series of subsequent impacts. Conventional honeycombs and lattice structures provide lightweight strength and stiffness, but they absorb mechanical impacts plastically and therefore require replacement after each impact event. One way to achieve elastic collapse and recovery is to fabricate lattice structures with extremely thin, hollow walls [1], [2].

One way to achieve elastic collapse and recovery is to fabricate lattice structures with extremely thin, hollow walls [1], [2]. Another strategy is to incorporate curved or buckled beams that exhibit negative stiffness or snap-through behavior upon application of a mechanical load. These negative stiffness honeycombs appear in several instantiations in the literature [3] - [6], including an early design presented by the authors and illustrated in Figure 1 [7], [8].Figure 1. Negative Stiffness Honeycomb [7]While the extruded two-dimensional honeycomb design in Figure 1 demonstrates nearly- ideal energy absorption (defined as mechanical energy absorption at constant force) under quasi-2170static and dynamic mechanical loads [7], [9], it is difficult to utilize it in a three-dimensional environment. It does not conform well to nonplanar, three-dimensional underlying surfaces such as cylinders, spheres, or other curvaceous objects.

It does not conform well to nonplanar, three-dimensional underlying surfaces such as cylinders, spheres, or other curvaceous objects. Also, it performs best when the loading is orthogonal to its base, placing the honeycomb in compression. Any shear loading on the honeycomb could cause the structure to bend out-of-plane and eventually yield and break. To address these challenges, a conformal negative stiffness honeycomb design was developed, fabricated, and experimentally tested.Conformal Negative Stiffness ElementThe conformal negative stiffness element is illustrated in Figure 2. The geometry of the conformal element in Figure 2a is constructed by generating the profile for a unit cell from the previous design (Figure 2b) and revolving it 360° about its central axis, as shown in the cross- sectional view in Figure 2c. Then, four corners of the element are removed, as shown in the cross-sectional view in Figure 2d to form the conformal element in Figure 2a.

Then, four corners of the element are removed, as shown in the cross-sectional view in Figure 2d to form the conformal element in Figure 2a. The corners are removed for two reasons: (1) to facilitate powder removal when fabricating the elements with polymer or metal powder bed fusion (PBF), and (2) to allow the elements to be tiled conformally across a surface, as illustrated in Figure 3. As shown, tiling allows the elements to protect complex curved surfaces from impacts.(a)(c)Figure 2. Conformal negative stiffness element (a) formed by revolving a unit cell cross section (b) about it central axis (c) and then removing four corners (d).2171(b)(d)Figure 3. Conformal Elements Arranged in a Tiled PatternConformal elements are designed to be fabricated with polymer and metal PBF in nylon 11 and 17-4PH stainless steel, respectively. Dimensions for the nylon 11 specimen are recorded in Figure 4.

Dimensions for the nylon 11 specimen are recorded in Figure 4. The curved shape of the beams is governed by the first mode shape of a buckled beam [10]. The dimensions of the beams are selected to achieve a prescribed force threshold, to prevent yielding of the beams, and to enable or prevent bistability, depending on the application.Critical Dimension Value (mm)Beam Thickness (t)1.66Center Thickness (ct)1.66Beam Length (L)50.0Bistability (Q)2.71Beam Height (h)4.50 = Q * tBeam Offset (off)2.08 = 1.25 * tBumper Width (buw)5.00Stem Width (sw)6.00Figure 4. Section view and Dimensions of Nylon Design2172Table 1.

Section view and Dimensions of Nylon Design2172Table 1. Selective Laser Sintered Nylon 11 Material Specifications, Obtained from a Tensile Bar Located Near the Specimens During the Build.Material SpecificationsValueYoung’s Modulus (E)1.21 GPaYield Strength (y)30 MPa (approximate)Yield Strain (emax)0.10 (approximate)The behavior of the beams is controlled by varying the thickness (t), length (L), andheight (h), of the beams, as indicated in Figure 4. Increasing the thickness or height of the beam results in a higher force threshold, meaning more force required to snap the beam from one buckled shape to another, as described in [7], [8]. Increasing the length of the beam results in a lower force threshold, and reduces the maximum strain in the beam. According to Qiu et al.

Theoretically, beams with Q values below 2.31 are monostable and2173(cid:4666)1(cid:4667)(cid:4666)2(cid:4667)(cid:4666)3(cid:4667)(cid:4666)4(cid:4667)(cid:4666)5(cid:4667)return to their original state after the load is released. With Q values above 2.31, the beam is stable in two configurations (i.e., bistable) and requires an external force to return it to its original configuration. Bistability is not desirable in this application, because the structure would require external intervention to protect against multiple impacts. However, the equations underlying the transition value for Q assume that the boundaries of the curved beam are perfectly rigid. Since the beams are held in place by a framework with some elasticity, higher Q values are possible without bistability. In previous applications, Q values as high as 2.71 were embodied in nylon 11 (polymer PBF) versions of the honeycombs in Figure 1 without observing bistability.

In previous applications, Q values as high as 2.71 were embodied in nylon 11 (polymer PBF) versions of the honeycombs in Figure 1 without observing bistability. For the conformal specimens, the Q value is varied from 2.00 to 2.71 by adjusting the beam height.The bumper width is set to a thickness that will not bend significantly under lateral loading from the curved beams as they snap through from one configuration to another. A thickness of 5 mm has been shown to adequately prevent this from happening in finite element simulations. The stem width is set to a thickness that will not shear under quasi-static loading. It can be increased if off-axis impacts are expected.With the critical dimensions set, the force required to buckle the beams, or forcethreshold, can be estimated using equations 1, 2, and 3 [10]. Those equations are valid for standard curved beams, as illustrated in the original honeycomb design in Figure 1.

Those equations are valid for standard curved beams, as illustrated in the original honeycomb design in Figure 1. To obtain the force threshold for a unit cell, the force threshold for a single curved beam is multiplied by two to account for double (concentric) beams in each instantiation. (The concentric beams encourage the beams to transition from a first-mode buckled shape to the opposite first-mode buckled shape via a third-mode, rather than a second-mode, buckled shape, thereby increasing the force threshold of the curved beams.) The same equations can be used to provide an estimate of the conformal design’s force threshold, if it is assumed that the conformal design is equivalent to two intersecting unit cells of the original honeycomb design, as represented by the two intersecting rectangles in a top view in Figure 5.Figure 5.

The Nylon Conformal Design Can Be Approximated as Two Intersecting Unit cells, as Shown in a Top View.Dimensions for the 17-4PH stainless steel specimen are recorded in Figure 6. The material properties for 17-4PH Stainless Steel, fabricated with direct metal laser sintering (DMLS) are recorded in Table 2, as provided by the manufacturer for the production process described in the next section. Of the available DMLS materials, 17-4PH stainless steel was2174selected for its low cost and high yield strain.

Section View of Metal DesignTable 2 Additive Stainless Steel 17-4PH Material SpecificationsMaterial SpecificationsValueYoung’s Modulus (E)170 GPaYield Strength (y)1241 MPaSafety Factor (sf)1.11Yield Strain (emax)0.0085From past DMLS projects, it was known that features as thin as 0.5 mm could be reliably sintered. This dimension was selected as the beam thickness to maximize the height of the beam (and therefore the force threshold) without exceeding the allowable yield strain (according to Equations 4 and 6). With the beam thickness selected, the length of the beam was set to 55.5 mm to maximize the length of the beams without interfering with holes for mounting the specimen to testing equipment, as shown in Figure 7.2175(cid:4666)6(cid:4667)BumperCenter BeamMounting HolesFigure 7. Metal Conformal Negative Stiffness ElementThe center thickness was increased relative to the nylon designs to provide a stiffer bracefor the four bumpers.

Metal Conformal Negative Stiffness ElementThe center thickness was increased relative to the nylon designs to provide a stiffer bracefor the four bumpers. FEA indicated that a center thickness equivalent to twice the thickness of the beams was large enough to prevent excessive lateral expansion of the bumpers. The offset thickness, bumper width, and stem width were set with criteria similar to the nylon design.Manufacturing MethodsThe nylon parts were selective laser sintered in PA D80-ST Nylon 11 from AdvancedLaser Materials on a 3D Systems Sinterstation HiQ-HS . The layer thickness was set at 0.004 in, the laser power at 40.0 W, and the part bed temperature at 182.5 °C.Figure 8. Selective Laser Sintering ProcessThe metal parts were manufactured on an EOS M270, with a layer thickness of 40microns, using the standard EOS 17-4PH build settings. Four metal specimens in total were manufactured. A sintered specimen is shown in Figure 9.

Four metal specimens in total were manufactured. A sintered specimen is shown in Figure 9. The specimens were built on their sides, as shown in Figure 10, in order to minimize the amount of support material needed. Nominally the overhang angle is 45 degrees, which is producible without support material in DMLS.2176After the sintering process, the entire build plate was heated to a temperature of 1900 °Ffor an hour and then air cooled. This process reduces residual stresses in the part and helps prevent warping as parts are removed from the build plate. After the specimens were removed from the build plate, holes for mounting an accelerometer, and mounting the specimen to the drop table, were machined.Figure 9. Sintered SpecimenFigure 10. Build OrientationAfter the machining process, the specimens were subjected to Hot Isostatic Pressing(HIP) to minimize residual porosity in the part.

Sintered SpecimenFigure 10. Build OrientationAfter the machining process, the specimens were subjected to Hot Isostatic Pressing(HIP) to minimize residual porosity in the part. The HIP process subjects the part to 2125 °F, at 14.75 kpsi, for a nominal time of 4 hours. After the HIP process, the specimens were then hardened using the standard H1150 heat treat process, which requires heating to 1150 °F, for 4 hours and then air cooling. During the HIP and heat treat processes, the parts were oriented as shown in Figure 11. This allowed the parts to sag under their own weight, disrupting the intended geometry of the beams. Figure 11 shows specimen #1, which suffered the most damage from the processes. One specimen was heat treated separately, oriented on its side as in Figure 10, and did not suffer from sag; its performance is compared to the damaged specimens in the experimental results section.2177Figure 11.

Heat Treat DamageFinite Element AnalysisFinite Element Analysis (FEA) was conducted using Abaqus/Explicit 2017. Acombination of solid and shell elements were used to create the finite element mesh. A part with the shell thickness rendered, is shown in Figure 12. The Abaqus S4R shell element, with enhanced hourglass control, was used to model the curved beams. The S4R element, is a quadrilateral shell with reduced integration. The Abaqus C3D8R element, with enhanced hourglass control, was use to mesh the remaining parts of the model, including the bumpers.Figure 12. Shell / Solid Combination MeshControlled displacement simulations were conducted to simulate the quasi-static tests anddynamic impact tests. For quasi-static simulations, a rigid body compressed the top of the specimen at a controlled rate, while the bottom of the specimen was minimally constrained.

For quasi-static simulations, a rigid body compressed the top of the specimen at a controlled rate, while the bottom of the specimen was minimally constrained. For the impulse simulations, the bottom nodes of the specimen were constrained using a prescribed node velocity.2178Quasi-Static TestingThe first test to be conducted on the specimens was a quasi-static test. The test wasconducted on a MTS Load Frame with a 100 kN load cell and fixed compression platens, at a rate of 2 mm/min. Figure 13 shows the unfiltered data from the test. There are no large spikes in force at maximum displacement because the MTS was programmed to begin unloading the specimen immediately before it fully compressed.Figure 13. Nylon Quasi-static TestFour different types of specimens were manufactured, with differing Q values. With a Qvalue of 2.00, the force threshold is low, shows little negative stiffness, and has a large flat region of zero stiffness.

With a Qvalue of 2.00, the force threshold is low, shows little negative stiffness, and has a large flat region of zero stiffness. The energy absorbed, (area between the load and unload curve), is also low. As the Q value is increased to 2.31, the force threshold becomes much higher, and two humps become visible in the plot, one for each row “snapping” through. Much more energy is also absorbed. As the Q value is increased further, the humps become more pronounced, and the energy absorption is also increased. If specimens were built with a Q value greater than 2.71, they would likely be bistable. For a bistable specimen, as one of the beams snapped through, the force would become negative.The force thresholds obtained experimentally are compared to those obtained from FEAand analytical approximations in Table 3. For the specimen with a Q value of 2.51, the FEA yielded a force threshold of 549.7 N, which is very close to the actual force threshold of 541.1 N.

For the specimen with a Q value of 2.51, the FEA yielded a force threshold of 549.7 N, which is very close to the actual force threshold of 541.1 N. Using the two column analytical approximation outlined in Figure 5 yields a force threshold of 543.5 N, which is also very close to the actual force threshold. For the Q value of 2.71, the force threshold estimated using the analytical method is 663.2 N, vs 535.3 N actual, while the FEA predicted 538.0 N. The analytical method assumes rigid boundary conditions for the beams, which can over predict the force threshold, especially for high Q values.2179Table 3 Nylon Quasi-Static Test ComparisonBistability Parameter2.512.71Quasi-Static Test541.1 N535.3 NFinite Element Analysis549.7 N538.0 NAnalytical543.5 N663.2 NA simple quasi-static test was conducted on the metal specimens at a rate of 5mm/min.Figure 14 shows the results of the first test cycle for the three damaged specimens.

Specimen #1 had the most damage and was tested first. The beginning of the test for specimen #1, shows a linear portion, which is caused by the compression platens correcting the slope of the top platform from a damage-induced slant to a horizontal position. The compression platens were held fixed to ensure that the top and bottom of the specimens were nominally parallel. The snap through of the first specimen was so loud that the technician stopped the test, as he thought the specimen had shattered. The test was then restarted, which explains the unload, load pattern for the first specimen. All four of the specimens exhibited a loud snap through sound for the first test. On all four specimens, the bottom beam was the first to snap through, represented by the first substantial decrease in force. After the first row snapped through, the force climbs again as the top beam is loaded.

After the first row snapped through, the force climbs again as the top beam is loaded. The top beam snaps through similarly to the bottom beam, but not as abrupt, and with less of a snapping sound. The second snap through is represented by the second substantial decrease in force. With both beams snapped through, the load is shunted to the bumpers causing the force to spike and the test to be terminated.All four of the specimens exhibited bistable behavior under quasi-static conditions, whichwas expected because of the very high Q value of the design. The thickness of the center beam also helped maintain the rigidity of the beam boundary conditions, contributing to the bistable response.Figure 14. Metal Quasi-Static Test2180All of the specimens were tested multiple times to investigate the repeatability of theirforce-displacement behavior. All three tests for specimen #2 are shown in Figure 15.

All three tests for specimen #2 are shown in Figure 15. The pronounced reduction in force threshold for the beams is due to plastic deformation, most likely where the beams attach to the bumpers. This area hardens however, leading to very repeatable results in the second and third tests.Figure 15. Specimen #2 Quasi-static TestsThe results of quasi-static testing of an undamaged Specimen #4 are compared with FEA in Figure 16. Due to an error by the manufacturer, specimen #4 was not included in first batch of parts sent to HIP, so it did not undergo the sagging deformation observed in the other specimens. It was heat treated to H1150 without HIP while oriented on its side, as in Figure 10. This prevented the damage seen in the other specimens. Due to its undamaged configuration, its initial force-displacement behavior is more linear than the damaged specimens. However, the force threshold remains similar, showing the robustness of the design.Figure 16.

However, the force threshold remains similar, showing the robustness of the design.Figure 16. Quasi-static Testing of Undamaged Specimen #4 vs. FEA2181The FEA in Figure 16 was conducted using material data from a DMLS build with thesame HIP and heat treatment process as the specimens. For the metal design a plastic model was also used in conjunction with the elastic model, in order to provide accurate results when the load approaches the yield strain. The Young’s Modulus of the material was 191.5 GPa, with a 0.2% yield stress of 839.5 MPa. This is very different from the manufacturer provided values in Table 2, and results in a yield strain of 0.00638, which is far lower than the yield strain for which the beams were designed. This is likely the reason for the large amounts of plastic deformation seen in Figure 15.The FEA analysis closely represents the actual test, with a simulated force threshold of698.4 N versus an experimental force threshold of 706.3 N.

The exact load curve does not match, likely due to differences between the model and the actual specimen, and due to kinetic energy error of the explicit FEA analysis, which simulates the compression test at a rate much faster than the quasi-static rate for computational efficiency.Drop TestingDrop tests were conducted on the first three metal specimens, using two different machines of the same operating type. As shown in Figure 18, the machines have a large boxed steel frame, on which two cylindrical rails are bolted in the vertical position. A large carriage rides on these cylindrical rails, and can be raised to a maximum height of approximately 15 feet. The carriage is also bungee assisted, to allow for greater impact velocities. The specimen is typically bolted to the top of the carriage, and the load is transferred to the specimen when the carriage impacts a large metal reaction mass at the bottom of the rig.

The specimen is typically bolted to the top of the carriage, and the load is transferred to the specimen when the carriage impacts a large metal reaction mass at the bottom of the rig. Felt, paper, and other materials are used to program the duration of the impulse. The technician fine tunes the impulse by adjusting the drop height, and programming material. Figure 17 shows a 15,000 g haversine impulse, with a 0.1 ms duration. The impulse is measured via an accelerometer mounted on top of the carriage, and an accelerometer mounted to the top of the specimen. The difference between the two accelerometer outputs shows the relative effectiveness of the specimen.DurationFigure 17. 15,000 g Haversine Impulse2182High-Speed CameraFigure 18. Drop Test EquipmentFigure 19 shows the result of a severe impulse with a duration of 0.088 ms.

15,000 g Haversine Impulse2182High-Speed CameraFigure 18. Drop Test EquipmentFigure 19 shows the result of a severe impulse with a duration of 0.088 ms. All of thedata in Figure 19, was sampled at 2.5 MHz, and filtered using a 3rd order Butterworth filter, with a low pass cut off frequency of 10 kHz. The maximum acceleration measured from the top of the specimen is 758.3 g versus a peak carriage acceleration of 11,700 g. The design has the effect of reducing the peak acceleration but extending the duration of the impact event.Figure 19. Specimen #2, 11,700 g impulseFigure 20 shows the impulse response of specimen 3 for three tests with approximatelythe same input impulse. As shown, the specimen’s impulse response is very repeatable.2183Drop RigSpecimenProgramming MaterialLightsFigure 20.

As shown, the specimen’s impulse response is very repeatable.2183Drop RigSpecimenProgramming MaterialLightsFigure 20. Specimen #3, 7900 g impulseTo simulate this impulse in Abaqus/Explicit, the acceleration is integrated to provide aconstrained node velocity for the bottom nodes of the specimen. For the case of the 15,000 g, 0.1 ms impulse, the entire element is assigned a prescribed velocity of -7.355 m/s. As the simulation begins, the bottom nodes are decelerated using the constrained node velocity in Figure 21.Figure 21. Constrained Node VelocityFigure 22 compares the Abaqus/Explicit predictions versus the experimental performanceof specimen 2. The peak acceleration agrees very closely between experiment and simulation: 784.8 g experimental versus 758.3 simulated. Differences can be attributed to several assumptions that are made in the finite element model. The first assumption is that the structure has not been stressed before the simulation.

Differences can be attributed to several assumptions that are made in the finite element model. The first assumption is that the structure has not been stressed before the simulation. In reality, the specimen has already undergone quasi- static testing, which led to plastic deformation. This has the effect of making the specimen less resistant to severe impulses, which is observed in Figure 22 in the lower force threshold and the increase in time required to respond to the input impulse. Future work is being done to incorporate a deformed mesh and material state as an initial state for the impulse model. This will allow the designer to run a controlled displacement simulation first, and then input the2184plastically deformed results into the impulse simulation. The second assumption is the use of the simplified input impulse shown in Figure 17. In reality the acceleration does not return to zero after the event.

The second assumption is the use of the simplified input impulse shown in Figure 17. In reality the acceleration does not return to zero after the event. The carriage has some upward momentum after its impact with the reaction mass and before the brakes are engaged. In order to decrease the analysis time this behavior is not modeled, but could be included later if more accuracy is desired.Figure 22. Specimen #2 vs. Abaqus SimulationThe first three metal specimens also exhibited damage from the heat treat. The simulation model did not account for this damage. The damage did not appear to have a significant effect on the dynamic response of the parts, but the damage did prevent testing at higher acceleration levels. Originally, the specimens were to be tested at a maximum level of 15,000 g, but on several 10,000 g tests, the base of the honeycomb descended at an angle, causing a large spike in acceleration due to the shunting of the load.

This phenomenon is shown in Figure 23, and it is likely caused by the sagging and warping of the curved beams. Increasing the load would have resulted in more acceleration spikes, which are not ideal for impact mitigation. It is suspected that specimen #4, would not have descended at an angle, allowing it to be tested at higher g levels.Load ShuntingFigure 23. Load Shunting on 10,000 g Test2185ConclusionThe conformal metal negative stiffness design shows much promise as an impactmitigation tool. The ability to tile the design, allows engineers to protect complex three dimensional surfaces against impacts from multiple directions. Since the design is made of stainless steel, it is resistant to heat, and can be used in unforgiving environments.

Since the design is made of stainless steel, it is resistant to heat, and can be used in unforgiving environments. Very severe impacts in the thousands of g’s can be reduced by an order of magnitude, which allows system engineers more freedom in placing sensitive components.AcknowledgmentsThis research was supported by an LDRD project from Sandia National Laboratories,entitled, “Additively Manufactured Shock Absorbing Engineered Materials.” Testing would have been impossible without the support of Sandia National Laboratories in Albuquerque, New Mexico. We would like to thank Tommy Woodall, Nicholas Leathe, and Audrey Morris-Eckart, for their help in scheduling testing, and providing design requirements. We would also like to thank Matthew Spletzer, and all the technicians and engineers at the Shock Lab, who spent weeks drop testing specimens, and providing subject matter expert support.2186References[1] T. A. Schaedler, A. J. Jacobsen, A. Torrents, A. E. Sorensen, J.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceADDITIVELY MANUFACTURED FASTENERSAmber Dressler1 2, William Scharnberg1, Anthony Abousleiman1, Samantha Harris1, and Phil New21Department of Mechanical Engineering, The University of Texas at Austin 2Sandia National LaboratoriesAbstractAdditive manufacturing (AM) has reached a critical point which enables production of complex, high resolution, custom parts from robust materials. However, traditional fasteners are still use to join these complex parts together. Integrating fasteners into additively manufactured parts is beneficial for part production but there is uncertainty in their design. To understand how the fasteners fit and function, mechanical property data was collected on the prototypes.

To understand how the fasteners fit and function, mechanical property data was collected on the prototypes. This data along with insights gained while building and testing the prototypes increased the knowledge base of design for additive manufacturing and build-to-build variability in selective laser sintering (SLS).IntroductionEngineers invented additive manufacturing in an attempt to reduce cost and time constraints binding the prototyping process. The project sponsor, Sandia, already benefits from this rapid, inexpensive prototyping in their broad array of research projects. Additive manufacturing (AM) has reached a critical point which enables production of complex, high resolution, custom parts from robust materials. AM also provides a newfound design freedom, allowing design focus to shift from how a part will be made to how a part needs to function. This remarkable shedding of existing design constraints paves the way for innovation of newly designed fasteners.

This remarkable shedding of existing design constraints paves the way for innovation of newly designed fasteners. The primary project aim was to develop fasteners that could be integrated into AM parts to alleviate the need for traditional fasteners. Since variability in part properties limits functional use of additively manufactured products, the secondary project aim was to increase Sandia’s understanding of build-to-build variability in selective laser sintering (SLS). Data was collected on build-to-build property variability, by building the fasteners in multiple build locations on the same machine over several days.Concept GenerationBefore creating design concepts, background information on SLS was collected to understand the process and its capabilities [1]. Existing design guidelines were used to help the design process, indicating what feature sizes would build successfully [2].

Existing design guidelines were used to help the design process, indicating what feature sizes would build successfully [2]. Understanding that there is variation in SLS part properties impacted design decisions attempting to and accommodate variation [3].Permanent Cap FastenerFigure 1 shows the initial SolidWorks model of the permanent pin and cap fastenerdesign; the cap is the upper part and the pin is the lower part.1791Figure 1. SolidWorks model of permanent cap fastenerThe cap contains four cantilevered beams each at 45° from the base that control the pins motion. Once the pin slides into the cap, the pin cannot retract because the beams prohibit motion in that direction. Figures 2 and 3 show cross-sections of the pin and cap in the unfastened and the fastened positions, respectively.Figure 2. Unfastened permanent pin and capFigure 3.

Figures 2 and 3 show cross-sections of the pin and cap in the unfastened and the fastened positions, respectively.Figure 2. Unfastened permanent pin and capFigure 3. Fastened permanent pin and capAs seen in Figure 1, the cap is not enclosed at the top to allow for the pin to be used multiple times for experimentation. In the final design, the cap would not be open at the top, making it a permanent fastener.1792Reusable Cap FastenerSince Sandia wanted both permanent and reusable fasteners, a variation of the pin and cap fastener that could be unfastened was designed. To do this, the AM capability of creating interior channels that would lock the pin in place was used. Figures 4 and 5 show SolidWorks models of the reusable pin and cap fastener in the unfastened and fastened positions, respectively.Figure 4. Unfastened reusable pin and capFigure 5.

Figures 4 and 5 show SolidWorks models of the reusable pin and cap fastener in the unfastened and fastened positions, respectively.Figure 4. Unfastened reusable pin and capFigure 5. Fastened reusable pin and capThe cap has multiple layers that the pin navigates before reaching its final fastenedposition, see the layers in Figure 14.Figure 6. Layers in the reusable cap fastener from bottom (left) to top (right)The pin inserts the cap through the bottom, or first, layer and moves up to the third layer. In the third layer, the pin can rotate 90° and retract down to the second layer. In the second layer, the pin cannot move while it is in tension. Once the tension is released, the pin can be removed from the cap by reversing the steps taken to fasten it.1793Permanent QuatrefoilFigure 7 shows the initial SolidWorks model of the permanent quatrefoil fastener design;the cap is the upper part and the pin is the lower part in Figure 7.Figure 7.

SolidWorks model of the permanent quatrefoil fastenerThis cap uses the corners of a quatrefoil as modified cantilever beams, and is inspired by the permanent cap fastener. Figure 8 shows the fastened cap and pin. The groove in the pin is designed to fix the cap in place.Figure 8. Fastened permanent quatrefoilBased on order of magnitude information gleaned from SolidWorks deformation studies of other fasteners, these quatrefoil corners were determined to be too thick for elastic deformation necessary to allow the pin to slide into the cap. The cap was therefore modified to allow the corners to function as the ends of four true cantilever beams, as shown in Figure 9.Figure 9.

The cap was therefore modified to allow the corners to function as the ends of four true cantilever beams, as shown in Figure 9.Figure 9. Modified permanent quatrefoil capFigure 9 also shows how the edges of the cap were extruded to facilitate a uniformapplication of force during fit testing.1794Guided Icosafoil FastenerFigure 10 shows the initial SolidWorks model of the guided icosafoil fastener design; inthis figure the cap is the upper part and the pin is the lower part.Figure 10. SolidWorks model of guided icosafoil fastenerThis fastener builds on the concept explored in the permanent quatrefoil, and its cap intends to use the filleted corners of 20 radially distributed circular arc sections in an icosafoil to mimic the function of cantilevered beams. This fastener further seeks to prevent rotation about the guiding axis of the pin by adding grooves to the pin which match respective grooves in the cap’s filleted icosafoil geometry.

This fastener further seeks to prevent rotation about the guiding axis of the pin by adding grooves to the pin which match respective grooves in the cap’s filleted icosafoil geometry. Figure 11 shows the guided icosafoil fastened.Figure 11. Fastened guided icosafoilFor similar reasons as with the permanent quatrefoil, additional ductility was introduced into the guided icosafoil cap, again by extruding cuts to create a series of miniature cantilever beam structures. Figure 12 presents the final cap for the guided icosafoil fastener.Figure 12. Modified guided icosafoil capThe ends of the beam structures are still textured to fit in the grooves on the pin, and withadded ductility this fastener is more likely to pass the fit test.1795Reusable Claw FastenerFigure 13 shows the initial SolidWorks model of the reusable claw fastener design; in thisfigure the claw is the lower part and the pin is the upper part.Figure 13.

SolidWorks model of reusable claw fastenerThis fastener builds on the use of cantilevered beams in order to maximize ductility, and therefore ease of fastening. In this initial design, both the claw and the pin employ cantilever beam structures which should elastically deform and then return to original geometry, interfacing as they fasten. After preliminary design work, it was noticed upon inspection that this fastener would disengage with any rotation of the pin about its vertical axis. In order to alleviate this concern, the cantilever beams of the pin were transformed into a cone, as presented in Figure 14.Figure 14. Fastened reusable claw with cone pinThis conical pin is less ductile than the pin having four cantilever beams, but is not subject to disengage with rotation about its vertical axis, being symmetrical about that axis. Additional material was added to the tip of the cone to create a more substantial connection as shown in Figure 15.1796Figure 15.

Additional material was added to the tip of the cone to create a more substantial connection as shown in Figure 15.1796Figure 15. Reusable claw with modified pinsPin modifications served to adequately fortify the junction without compromising excess ductility. This fastener also contains the potential to be tested for failure not only in tension, but also in compression, which is another area of interest for Sandia.Puzzle Piece FastenerThe SolidWorks model of the puzzle piece fastener is shown in Figure 24 and Figure 25. The part to the right in Figure 24 will henceforth be referred to as Part 1, and the part to the left in Figure 24 will be called Part 2.Figure 16. Disassembled puzzle piece fastenerFigure 17. Assembled puzzle piece fastenerThis fastener functions by taking Part 1 and sliding it into the corresponding slots in Part 2.

Disassembled puzzle piece fastenerFigure 17. Assembled puzzle piece fastenerThis fastener functions by taking Part 1 and sliding it into the corresponding slots in Part 2. Once the two pieces are attached as in Figure 25, the user would then pull upward on Part 1 to lock it in grooves present in the slots to lock the parts together.1797Locking Shaft FastenerThe locking shaft fastener attaches two rods together, end to end. Figures 18 and 19 show the two parts of the fastener. This fastener functions by taking Part 1 in Figure 18 and sliding it into Part 2 in Figure 19.Figure 18. Locking shaft fastener part 1Figure 19. Section view of locking shaft fastener part 2Propagating Claw FastenerFigure 20 shows the propagating claw fastener, designed by analogy with the caps of adry-erase marker.Figure 20. SolidWorks model of propagating claw fastenerThe design can be divided at its neck into a pin on top and clasps below.

SolidWorks model of propagating claw fastenerThe design can be divided at its neck into a pin on top and clasps below. The clasps form a claw which receives the pin by flexing outwards as a force is applied, deflecting enough to allow the pin through. Once the pin has entered, the claw returns to its undeflected position and, with the extruded lip on each clasp acting as an abutment surface, prevents the pin from retracting. Figure 21 shows the fastener in use, with the clasps encapsulating the pin.1798Figure 21. Attached propagating claw fastenerReusable Clip FastenerIn a bike helmet, the clip can potentially unfasten if it receives a force from the side. For this design, the housing for the clip was made so the clip cannot be removed unless there is an active attempt to unfasten it with a tool. Figure 22 shows a SolidWorks model of the housing (left) and the clip (right).Figure 22.

Figure 22 shows a SolidWorks model of the housing (left) and the clip (right).Figure 22. SolidWorks model of the reusable clip fastenerThe top and bottom members of the clip deflect towards the center member when the clip is inserted into the housing. The width of the cavity within the housing expands near the center of the housing; once the clip reaches the expansion, the deflected members snap back to their original position to fasten the clip. Figure 23 shows the clip in the fastened position.Figure 23. Fastened model of the reusable clip fastenerThe slot in the housing allows for a tool to unfasten the clip by deflecting the top and bottom members towards the center. After the members are deflected, the tool can slide the clip out of the fastener.1799PrototypesTo test functionality all nine fastener designs were prototyped, see Figure 24, in nylon 12on the SLS machine at UT-Austin [4].Figure 24.

Prototypes of fastener conceptsUpon receipt of the prototypes, fit tests were conducted to determine if the designs fastened as expected. Out of the nine fastener designs, six fastened as expected and three did not fasten. Two of the three designs which failed the fit test used flexible members as their fastening mechanisms and would have benefitted from iterative prototyping. The permanent pin and cap design did not fasten because the cantilever beams around the cap were more brittle than expected; the beams snapped off before the pin was fastened. Similarly, flexibility issues caused failure in the guided icosafoil prototype. The locking shafts fastener failed the fit testing due to tolerance issues. The remaining designs were functional; however, certain drawbacks were realized during the fit tests. The reusable pin and cap allowed for excessive shifting of the pin within the cap.

The remaining designs were functional; however, certain drawbacks were realized during the fit tests. The reusable pin and cap allowed for excessive shifting of the pin within the cap. The permanent quatrefoil appeared to be fairly weak and was unable to remain fastened against even small forces.Design Selection and ModificationIn order to down select to the final designs, the functionality and testability of each design were considered. First, the permanent pin and cap, the guided icosafoil, and the locking shafts were eliminated due to their lack of functionality. Next, the permanent quatrefoil and the reusable claw were eliminated due to their relatively low strengths. Then, the puzzle piece fastener was eliminated due to testability; a large amount of material would be needed to attach sections for the tensile testing clamps.

Then, the puzzle piece fastener was eliminated due to testability; a large amount of material would be needed to attach sections for the tensile testing clamps. Finally, the reusable pin and cap, the reusable clip, and the propagating claw were chosen for more testing.After design selection, design modifications were made to the chosen fasteners to accommodate tensile testing. Tapered extrusions were added on each end of the fastener to fit into the tensile testing machine. Stress analyses were conducted in SolidWorks to ensure the sections did not add stress concentrators to the design. The results of the tensile extrusion simulations are presented in Figures 25-27 below.1800Figure 25. Stress analysis of design modification of the propagating claw fastenerAs expected, the largest stress concentrations shown in Figure 25 are in the ‘neck’ of the design.

Stress analysis of design modification of the propagating claw fastenerAs expected, the largest stress concentrations shown in Figure 25 are in the ‘neck’ of the design. The results of the simulations also indicated stress concentrators at a joint between two extrusions of equal cross section on the propagating claw fastener. These were experimentally determined to be software artifacts stemming from the extrusion process rather than realistic indicators of existing stress concentrators.Figure 26. Stress analysis of design modification of the pin and cap fastenerThe results in Figure 26 showed that the reusable pin and cap would fail at the abutmentsurfaces and not at the additional extrusions as desired.Figure 27. Stress analysis of design modification of the clip fastenerSimulations presented in Figure 27 also demonstrated that the reusable clip would function as expected.

Stress analysis of design modification of the clip fastenerSimulations presented in Figure 27 also demonstrated that the reusable clip would function as expected. Bending occurs in the clip housing simulation as a result of fixing one end of the clip while applying a force to the other, which is the required method for finite element1801tensile simulations in SolidWorks. Bending in this case is caused by the design’s asymmetry due to its window.ExperimentsThe build-to-build variation testing was separated into two categories: location-to- location and day-to-day. For location variation, each fastener was built in nine locations on the build platform. For day variation, the location variation test was built three times. A tensile bar was placed in each group of fasteners for comparing the strength of the fastener with the strength of the AM material. The build layout for the variation testing is shown below in Figure 28.123456789Figure 28.

The build layout for the variation testing is shown below in Figure 28.123456789Figure 28. Build layout and numbering for variability testingA SLS machine at Stratasys Direct Manufacturing using glass-filled nylon 12 was used for the variation testing [5]. A single machine was used for all three builds to eliminate a variable from the variation testing. Between the first and second build, the SLS machine underwent maintenance due to a failed build.ResultsTensile tests were completed on the tensile bars and the fasteners from the first build; the results are shown below in Figure 29. Since the fasteners have varying cross-sectional areas, maximum load was used as the output metric for comparable results.

Since the fasteners have varying cross-sectional areas, maximum load was used as the output metric for comparable results. The results for the propagating claw were not included because one of the legs fractured before the test was conducted.18021800160014001200)N( d a o L x a M1000800Tensile BarsPin and CapClip600Propagating Claw40020001235 4 Build Location6789Figure 29. Tensile test results from the first buildThe reusable pin and cap design exhibited the closest maximum loads in comparison to the tensile bars with the reusable clip being the furthest from them. The reusable clip failed at lower loads, in part, due to having the smallest abutment area. All of the specimens presented variation in maximum load dependent on the location.

The reusable clip failed at lower loads, in part, due to having the smallest abutment area. All of the specimens presented variation in maximum load dependent on the location. With this number of samples, a clear trend in the variation cannot be determined.Once all three builds were tested, the maximum load data for each design was graphed to compare location data and day data. These graphs are presented in figures 30-32. After the maintenance between the first two builds, the propagating claw experienced tolerance issues that prohibited it from fastening; therefore, maximum load data for the propagating claw fastener was not collected for the final two builds.1803200018001600)1400N( d a o L x a M12001000800Build 1Build 2600Build 340020001234 6 5 Build Location789Figure 30.

Max load of each tensile bar from the variation testingIn builds 1-3, the average maximum loads withstood by the tensile bars were 1595 N, 795 N, and 900 N, respectively. Figure 30 shows that there is a significant difference between the results of builds 1 and 2 while builds 2 and 3 are more similar. The only difference in builds was the maintenance conducted between builds 1 and 2.12001000)N800( d a o L x a M600400Build 1Build 2Build 320001234 6 5 Build Location789Figure 31. Max load of the reusable pin and cap design from the variation testingIn builds 1-3, the average maximum loads withstood by the reusable pin and cap were 954 N, 671 N, and 701 N, respectively. Figure 31 shows the similar variation between pre- maintenance and post-maintenance builds.1804250200)N( d a o L x a M150100Build 1Build 2Build 35001234 6 5 Build Location789Figure 32.

Max load of the reusable clip design from the variation testingIn builds 1-3, the average maximum loads withstood by the reusable clip were 226 N, 199 N, and 165 N, respectively. The reusable clip did not experience the same trends in variation between pre-maintenance and post-maintenance builds.While variation was present between all locations and days, the machine maintenance seems to have had an impact on the maximum load reached during the tensile tests. At this point, more samples are needed to have a better understanding of variation in SLS part properties.Future WorkIterative design studies of each of the fastener concepts would greatly enhance their functionality and strength. Information on how the fasteners perform at a range of different sizes would provide beneficial information for their integration into AM parts. Studies in additional materials are necessary to increase the impact of these fasteners on AM design.

Studies in additional materials are necessary to increase the impact of these fasteners on AM design. An increased sample size will provide a better representation of variation across the build plate. With more samples, a property map of the build plate would be beneficial to AM part production.ConclusionsThis paper describes nine fastener concepts that can be integrated into additively manufactured parts to reduce the need of traditional fasteners. Out of the fasteners tested, the reusable pin and cap withstood the highest maximum load. The strength of the reusable clip could be improved with a redesign to have larger contact area between its abutment surfaces. Adjusting the tolerance in the propagating claw fastener is necessary to account for dimensional variation between builds.

Adjusting the tolerance in the propagating claw fastener is necessary to account for dimensional variation between builds. For the remaining six designs, iterative design would greatly improve their functionality and feasibility as AM fasteners.In addition to testing functionality, replicating the fasteners showed that there is variation in part strength from location-to-location and day-to-day. A potential cause of the large1805difference between the first build and the remaining builds was that the SLS machine underwent maintenance due to a failed build between builds 1 and 2. More samples are needed to increase the understanding of build-to-build variability in selective laser sintering.AcknowledgementsWe wish to acknowledge Sandia National Laboratories for sponsoring this research, with many thanks extended to their technical staff who offered assistance and guidance. Special thanks are given to Dr.

Special thanks are given to Dr. Carolyn Seepersad, for her expert insight, mentorship, and generosity of time. SLS prototyping at UT Austin was facilitated by Mark Phillips. Lastly, we would like to thank Brad Pagano of Stratasys Direct Manufacturing, who worked closely with us to ensure our builds were completed to the required specifications.References[1] C. Seepersad, Powder-Based SFF Systems, The University of Texas at Austin, Department ofMechanical Engineering, 2015.[2] C. Seepersad, T. Govett, K. Kim, M. Lundin and D. Pinero, "A Designer's Guide forDimensioning and Tolerancing SLS Parts," The University of Texas at Austin, Department of Mechanical Engineering, 2012.[3] D. L. Bourell, T. J. Watt, D. k. Leigh and B. Fulcher, "Performance limitations in polymer laser sintering," in 8th International Conference on Photonic Technologies LANE 2014, 2014.[4] Stratasys Direct Manufacturing, "NyTek 1200 PA: Laser Sintering Material Specifications,"2015. [Online].

ditively Manufactured Flexible Fluidic Actuators For Precision Control in Surgical ApplicationsJonathon E. Slightam & Vito R. Gervasi Rapid Prototyping Research, Milwaukee School of Engineering, 1025 N. Broadway, Milwaukee, WI 53202AbstractPreviously, the Milwaukee School of Engineering (MSOE) demonstrated a dexterous tele-operational robotic system where actuators, linkages were fabricated simultaneously using Selective Laser Sintered Nylon 12. Primary motivation for this research was to conceive novel fluid power actuators that were inherently safe, compact, and Magnetic Resonance Imaging (MRI) compatible for surgery and rehabilitation. Although the concept of fabricating MRI compatible fluid power devices was demonstrated, further proof of precision control was needed.

Although the concept of fabricating MRI compatible fluid power devices was demonstrated, further proof of precision control was needed. The design and implementation of additively manufactured flexible fluidic actuators (AMFFA) for precision control, best practices, and the comparison of these actuators with other actuation technologies are presented.joints, andIntroductionPurposeProject 2G, “Robotic Surgery and Rehabilitation via Compact, Integrated Systems” is focused on the research and development of compact surgical and rehabilitation devices for use inside the confines of MRI equipment. Expanding the use of fluid power in medical applications is primary technological motivation for this work. It provides an efficient way to implement compact precise control in harsh environments such as the strong magnetic fields associated with MRI machines while also tapping into existing pneumatic power readily available in hospital infrastructure [1].

While MRI-compatible pneumatically powered actuators and robotic systems are seldom commercially available, those available aren’t inherently safe. The use of additive manufacturing (AM) technologies in fluid power and robotics has continued to illustrate significant technological benefits that typically yield greater performance through reductions in mass, size, and energy consumption. Using AM also has introduced the realization of simultaneous manufacturing, the process of fabricating systems as a single entity opposed to traditional methods that assemble separate components after they’re produced. Simultaneous manufacturing allows for actuators, mechanical structures, mechanisms, and sensors to be fabricated at the same time, producing a framework similar to what’s often illustrated in nature.

Simultaneous manufacturing allows for actuators, mechanical structures, mechanisms, and sensors to be fabricated at the same time, producing a framework similar to what’s often illustrated in nature. For robotic engineers, the ability to seamlessly manufacture robotic systems paralleling the fabrication of biological systems provides a tool to develop robots that are fundamentally what’s desired to be mirror images or extensions of ourselves. The purpose of this research is to blend the ideology behind the ability to use additive manufacturing to simultaneously manufacture robotic systems and technological need to develop inherently safe and compact robotic systems for MR-image guided surgical interventions.

This endeavor presents technical challenges in the use of AM, discusses modeling strategies for precision control, and most importantly presentsthe inherent safety in the design of an additively manufactured flexible fluidic actuator (FFA) for surgical applications.ScopeResearch and technical artifacts from MSOE’s involvement with the National Science Foundation’s Engineering Research Center for Compact and Efficient Fluid Power are presented herein. The scope encompasses similar research conducted by other research institutions and MSOE and medical motivation for the current work is presented.

The scope encompasses similar research conducted by other research institutions and MSOE and medical motivation for the current work is presented. The paradigm of simultaneous in simultaneous manufacturing design manufacturing, modeling strategies, development results and motivation for future investigations in addition to breakthrough applications are discussed.in development of FFAs,technical challengesBackgroundResearch by Delaurentis, Mavroidis, and Won demonstrated the feasibility of using AM for simultaneously fabricating mechanisms using selective laser sintering (SLS) and stereolithography (SLA) [2-6]. However, only some electromagnetic actuators could be embedded during fabrication by pausing the fabrication process mid-build and is feasible only with SLA and fused deposition modeling (FDM). Festo Inc.

Most recently, the Automation Manufacturing and Robotics (ARM) group at Oak Ridge National Laboratory and MSOE outline guidelines for use in the design of additively manufacturedfluidically integrated robot systems and components for electron beam melting and polymer SLS processes, respectively [13,14]. Engineered fluid power robot components and devices have illustrated reductions in weight up to 80% and also up approximately 90% reduction in volume [13,14]. Yet, the most limiting technical barrier on smaller sizes is removal of unsintered and lightly sintered powder around and inside fabricated components [13,14]. This is due in part to the residual energy dissipated into the powder surrounding parts in the powder bed ultimately causing powder degradation [15,16].

This is due in part to the residual energy dissipated into the powder surrounding parts in the powder bed ultimately causing powder degradation [15,16]. Fluidic channels and internal recessed geometries out of direct line of sight make it difficult for powder removal and is sometimes the leading cause of an unacceptable design as Figure 2 illustrates for an embedded universal joint. Knowing distinct limitations in design that are highly repeatable are needed for the design of MRI-compatible actuators and robotic systems.Figure 2: Embedded universal joint depicting required force to shear unsintered powder.Project 2G’s primary focus for MRI guided interventions is hypothermal ablation for treatment of epilepsy. Epilepsy is known to effect approximately 50 million people worldwide, 40% of which are unresponsive to medication [17]. In addition, 7 – 17% of those diagnosed with epilepsy experience sudden death [17].

In addition, 7 – 17% of those diagnosed with epilepsy experience sudden death [17]. Of the 2 million Americans diagnosed, 50% remain untreated and surgical procedures have only illustrated a 70% success rate [18-20]. The use of MRI guidance allows for precise minimally invasive procedures to be executed by using imaging as feedback for closed loop position and temperature control [1]. Present MRI-compatible actuators are pneumatically powered 2-way piston cylinders fabricated with MRI-compatible materials. However, 2-way piston cylinders aren’t inherently safe for image guided interventions and surgical procedures. If a pressurized pneumatic supply or return line were to disconnect, a valve circuit were to short and fully open, the cylinder could fully retract or actuate the entire stroke length. This presents an immediate danger to uses in clinical trials and is unacceptable.

This presents an immediate danger to uses in clinical trials and is unacceptable. Clinical needs and technical deficiencies illustrate a strong need for inherently safe actuation devices operable in MRI environments.This research discusses the use of additive manufacturing to simultaneously manufacture inherently safe devices using flexible fluidic actuators for image guided interventions andsurgical procedures, current and previously proposed system models for FFAs, actuator performance, recommended approaches, and potential applications. The culmination of previous work as well as clinical and technical needs fully support the motivation for this applied research.Additively Manufactured Flexible Fluidic ActuatorsDesign and Approach ParadigmUse of additively manufactured FFAs allow for simultaneous fabrication of pneumatically powered robotic systems and devices.

As previously proposed, the class of actuator allows for freeform fabrication around or through specific geometries allowing for effective power transmission and reducing the overall system size. For designing a device that drives a hypodermic needle or an ablator tip for MRI-guided interventions, the design process illustrated in Figure 3 was used from prior work and implemented. It helps convey the overall ideology behind simultaneous manufacturing and that consolidation of multiple components can lead to more compact systems.Identification of need“Traditional Architecture”Definition of Problem“Simultaneous Manufacturing Architecture”SynthesisAnalysis and optimizationOptimization and Design for AMDesign for CleanoutPost ProcessingEvaluationPresentationHigh level designs integrating multiple aspects of machines into a single entity.

• Structural designs that conform to complex shapes to resuce volume or mass.Structural, thermal, and fluidic optimization by themselves and as multifunctional optimized components.Structures acting as a guide for a FFA should have ~ 0.5mm clearanceClean out holes generally need to be greater than 1-2 mm.Fluid passage entry ways need to be in direct line of sight for tools such as dental picks and dremel cleaning heads to enter.Joints should includemultiple perforation holes for cleanout and have clearances approximately greater than 0.35mm.Clean out holes need to be in direct line of sight for sealing via welding a filiment into the passage entry way. Thus, a soldering iron tip, or custom heating element should be allowed to reach it.certain surfaces mayrequire sealing depending on the fluid used.

Thus, a soldering iron tip, or custom heating element should be allowed to reach it.certain surfaces mayrequire sealing depending on the fluid used. Silicones typically work very well.Figure 3: Simultaneous manufacturing architecture and design outlines [21].Using the design process and guidelines specifically aimed at generating fluid powered robotic devices actuated with FFAs, devices that are extremely compact can be developed.Following these guidelines can produce repeatable results that are inherently more compact than conventionally manufactured counterparts because of sub-system & component integration that is part of the overall changing framework in robotic system and device design.Modeling of Additively Manufactured Flexible Fluidic ActuatorsPrior research using FFAs used a dynamical model that was linear in nature. However, nonlinear models were proposed by Slightam and Gervasi to properly describe the visco-elastic behavior of the polymer membranes [22].

However, nonlinear models were proposed by Slightam and Gervasi to properly describe the visco-elastic behavior of the polymer membranes [22]. These two representations are depicted in Figure 4 part a) and part b), respectively.η1ξ1η2ξ2ba)b)Figure 4: Tradition linearized model vs. previously proposed visco-elastic model.The use of a spring-mass-damper (SMD) model implies a straightforward approach because of its common use in controls, but requires use of nonlinear sliding mode control (SMC) to account for nonlinear characteristics as previously demonstrated [23]. A visco-elastic model was developed to accurately describe the physics of the dynamic system and determine whether or not it is better than the SMD model. One of the primary tradeoffs to consider is the time invested in system design and characterization, which is discussed in the results of implementing a new AM FFA.Performance ComparisonThe use of bellows as actuation is not relatively new.

However, recent implementation and new developments have arisen from needs for complaint actuators in human-machine interaction. It is because of this need for compliance that fluidically actuated bellows or FFAs have illustrated a strong reemergence. Qualitatively, FFAs are compliant when powered by hydraulic fluid and more so with compressed air but little known work has illustrate how recent developments of FFAs compare with other actuation technologies [23]. This comparison is an essential part in high level robotic system design and helps select the most suitable means of actuation for a particular actuation, while providing insight into other applications beyond the scope of current surgical and rehabilitation applications.Typically, strain, stress, specific power, bandwidth, and stiffness, and often efficiency are traits for comparing actuation technologies to one another [24].

Figure 5 depicts a generalized form of an actuator that helps define these traits for comparison.LsmF, WAFigure 5: Free body diagram of a general actuating element.Where s is the actuator stroke length and L the actuator’s overall length, F, is the produced force from the actuator, A, the cross-sectional area of the actuator, W, is the work out from the actuator, m, is the mass, and k is the actuator stiffness.The strain of an actuator is defined by the actuation stroke length per the total actuatorlength as Eq. (1) depicts.𝜀 =𝑠𝐿(1)Stress of an actuator can be characterized by the produced force per cross sectional area,defined by Eq. (2).𝜎 =𝐹𝐴(2)Specific power can be determined from the mechanical work produced from the actuatorper its total mass as Eq. (3) describes.𝜌𝑊 =𝑊𝑚(3)Bandwidth, also referred to as the response of a diaphragm, is defined by Eq.

(3) describes.𝜌𝑊 =𝑊𝑚(3)Bandwidth, also referred to as the response of a diaphragm, is defined by Eq. (4).1/2 ]10.21𝑔𝐷2𝜋𝑎 [ Where g is the gravitational constant, D is the flexural rigidity, and w is the specific weight of the material, where these material constraints affecting the diaphragm responses are more thoroughly described by Giovanni [25]. However, one of the limiting factors for response is the fluid power system and is on average approximately 50Hz. Since diaphragm responses are typically greater than this, it’s suffice to use this approximation for the bandwidth.𝑓 =ℎ𝑤(4)Lastly, the stiffness or load holding capabilities of an actuator can be determined by itsstiffness per unit length.𝐾 =𝑘𝐿Efficiency is regarded to be an important parameter for actuator performance comparison as well. It has been declared that the efficiency of FFAs is greater than those of their piston- cylinder counterparts [26].

It has been declared that the efficiency of FFAs is greater than those of their piston- cylinder counterparts [26]. However, this claim may be fallacious without proper validation as FFAs are also subjected to energy losses in the fluid, losses from stored strain energy, friction, and is also dependent on the work out of the actuator, that is dependent on the output force. The control system of FFAs is also a contributing factor to the overall efficiency. Use of high speed digital solenoid valves allows efficient discrete control of air in and out of the actuator. Otherwise, a less efficient approach is pressurization via backpressure in the FFA caused by restricted flow on the exhaust passage from an orifice.This speculation and criticism provides motivation for validating these claims and conducting non-dimensional performance analysis on AM FFAs to improve the understanding of the entire spectrum of capabilities.

However, validating claims or speculations of high efficiencies of FFAs are not included because of the technical depth is beyond the scope of this technical paper.ResultsAs a result of using the ideology of simultaneous manufacturing and guidelines outlinedin Figure 3 and the clinical need to develop and inherently MRI-compatible device, the linear stepper motor for needle actuation depicted in Figure 6 was developed.Clean Out HolesPinching MechanismToroidal BellowsOptical Encoder & MountSimulated AblatorTest FixtureFigure 6: MRI-compatible linear stepper motor for needle insertion.The inherent safety of the actuator is through its discrete sequential stepping design. For example, in the event of a system malfunction as previously discussed the actuation or retraction of the needle would be limited to the discrete step size. In addition, sub-step control is possible for precise needle placement between discrete step lengths.

In addition, sub-step control is possible for precise needle placement between discrete step lengths. The actuator implemented a US- Digital EM1 transmissive optical encoder module with 500 lines per inch.(5)Using SMC, precision control was achieved. Steady state error of 25 microns was illustrated during multiple tests such as the one depicted in Figure 7. Only 15 microns of precision could feasibly be achieved because of the digital encoder used.Figure 7: SMC Results.The performance of the actuator followed theoretical predictions of stiffness and near linear behavior on initial actuation as Figure 8 depicts. This also allowed for the hysteretic damping to be determined experimentally.Figure 8: Toroidal FFA position as a function of internal pressure.As expected, the actuator exhibited visco-elastic behavior that the controller was able to compensate.

This is especially apparent in Figure 9 and 10 that illustrate the exponential decay and growth of pressure needed to achieve the desired motion in Figure 7 and the modulus behavior, respectively.Figure 9: Pressure applied to FFA in SMC trial.Figure 10: Approximate Modulus of FFA during SMC trial.From experimental data using the SMC the visco-elastic model was once again tested and the results are depicted in Figure 11. However, it was only conducted for one actuation cycle as thevisco-elastic model is time dependent and employing a real-time system model that integrates the Boltzmann’s superposition principal is beyond the scope of this work [27].Figure 11: Visco-elastic dynamic model results.Given the actuator parameters, the performance was also characterized.

The pneumatically powered toroidal FFA was compared to other actuation technologies in Table 1.Table 1: Actuation technology comparison [11, 28, 29].This gives insight into potential applications outside of MRI compatible surgery and rehabilitation in addition to an initial technological assessment of the device. The device can be used in a modular fashion and eventually replace the original 5 degree of freedom cannula robot as conceptualized in Figure 12. The original implemented 5 degree of freedom cannula robot used a volume of approximately 2592 cubic inches (42,475 cubic cm), while it is projected that a modular 5 degree of freedom robot using FFAs could occupy a volume of 260 cubic inches(4,261 cubic cm).

Thus resulting in a projected volume reduction by as much as 90% using the methodologies applied to develop the device.Figure 12: Conceptual modular 5 DOF cannula robot.Potential applicationsBased on the empirical results of implementing the ideology of designing devices and systems with FFAs, other potential applications outside of MRI-compatible surgery can be identified. The precision demonstrated that medical applications outside of the one discussed herein can achieve, such as full surgery rather than interventions, in addition to rehabilitation. The similarities between human muscle and the actuator performance characteristics illustrate that rehabilitation devices such as orthoses and even prostheses could potentially be applicable, as FFAs have been used for in the past, although not with AM.

Branching out into other medical applications may provide for mass customization with orthoses.ConclusionsUse of the paradigm to simultaneously manufacture devices with integrated fluidic actuators and implementation of SMC has produced a device that is inherently safe because of its discrete motion, compact, and precise. While models used with the SMC don’t describe the true nature of Nylon 12 FFAs, it’s a more practical and efficient approach when considering the overall design and implementation process. Several steps in experimentation can be skipped that are needed to determine the visco-elastic constants in the MKV model, reducing the time to implement the device. Furthermore, the use of the MKV model has yet to illustrate precision greater than a robust SMC that demonstrates precision consistently near that of the resolution of the position sensors used.Initial assessments show that the performance of pneumatic FFAs can be similar to the traits of human muscle.

While past conjectures have stated that efficiency can be greater thanthat of piston cylinder counterparts, qualitative discussion of FFAs suggest that efficiencies could be lower than originally claimed. This analysis gives insight to how such devices should be properly designed in the future, to maximize both efficiency and work out of systems. It also identifies that such devices aren’t as promising as once thought, but still practical for many applications requiring compact devices that are also customizable.AcknowledgementsSpecial thanks are given to David B. Comber at Vanderbilt University for outstanding performance in collaboration with MSOE for his help in guiding the mechanical design as well as performing the SMC design. Thanks are also given to the National Science Foundation’s Engineering Research Center of Compact and Efficient Fluid Power for their continued support and funding of this work.References[1] Comber, D. B., Cardona, D., Webster III, R.

Wixom, MI 48393, USACorresponding author: shamsaei@auburn.edu Phone: (334) 844-4839AbstractIn this study, the effect of post-process heat treatment on the microstructure and mechanical properties of Hastelloy-X superalloy fabricated via two different additive manufacturing technologies, namely, laser beam powder bed fusion (LB-PBF) and laser powder directed energy deposition (LP-DED), is investigated. Microstructure was examined using scanning electron microscopy (SEM) and electron backscattered diffraction (EBSD) analysis, while mechanical properties were evaluated by macro- hardness testing using the Rockwell B method. Microstructure of the alloys was studied thoroughly after several heat treatments that involve stress-relieving (at 1066°C for 1.5 hours), hot isostatic pressing (HIP at 1163°C for 3 hours under 103 MPa pressure), and/or solution treatment (at 1177°C for 3 hours).

The results revealed that, for both LB-PBF and LP-DED Hastelloy-X, the post-process heat treatments resulted in uniform grain structure as well as partial dissolution of carbides, although they have different grain sizes.Keywords: Additive manufacturing, Hastelloy-X, microstructure, grain size, macro- hardness.IntroductionDue to Hastelloy X’s ability to retain high strength at elevated temperatures even in highly corrosive environments, it is extensively used to fabricate parts with a complex structure typically used in power generation gas turbines and aerospace industries [1,2]. However, as it is quite challenging to machine these Ni-base parts, additive manufacturing (AM)—having the capability to fabricate finished complex parts— can be a suitable alternative manufacturing process for this alloy. Among different AM methods, laser beam powder bed fusion (LB-PBF) and laser powder directed energy deposition (LP-DED) are two well-known and extensively used techniques.

Among different AM methods, laser beam powder bed fusion (LB-PBF) and laser powder directed energy deposition (LP-DED) are two well-known and extensively used techniques. In the LB-PBF process,923powders are spread uniformly over a build platform, and a heat source melts the powders together to fabricate parts [3]. In the LP-DED technique, powder particles are injected for deposition and melted to fabricate parts [4].In spite of AM techniques providing several advantages over the traditional manufacturing methods, the mechanical properties of AM parts may differ from those of the conventionally fabricated parts. For instance, the microstructure of AM parts often differs from the conventionally fabricated parts due to the unique thermal history during the manufacturing process leading to the process-induced volumetric defects and residual stresses, which in turn may impact the mechanical properties of the AM components.

The larger variation in characteristic thermal history among different AM technologies also exacerbates this. Hence, utilizing the same heat treatment procedures proposed in the literature for the wrought counterparts on the AM components may not necessarily result in similar mechanical performance [5]. Therefore, there is a knowledge gap for thorough verification of the applicability of the conventional heat treatment processes proposed for the wrought component on the AM materials.Hastelloy-X is a primarily solid solution and partially carbide strengthened Ni-base superalloy [6,7]. The absence of Nb in Hastelloy X, a potential δ and Laves phase former, may potentially give it better mechanical properties in as-fabricated conditions compared to other alloys with Nb.

The absence of Nb in Hastelloy X, a potential δ and Laves phase former, may potentially give it better mechanical properties in as-fabricated conditions compared to other alloys with Nb. Although microstructure and mechanical properties of Hastelloy X has been investigated in both wrought [2,6–8] and various AM conditions [9–14], the effect of multiple-step heat treatments, which are often necessary for AM parts, on both microstructure and mechanical properties of AM Hastelloy-X fabricated by different AM processes and their comparison is still lacking. In this paper, the effect of multiple-step heat treatments on the microstructure and mechanical properties of LB-PBF and LP-DED Hastelloy-X has been thoroughly investigated and compared.Experimental ProcedureCylindrical bars of Hastelloy-X with 14 mm diameter and 90 mm height were fabricated using both LB-PBF and LP-DED techniques.

The LB-PBF Hastelloy-X bars were fabricated by SLM280 2.0 machine using a pre-heated build platform (recommended by SLM solutions) at 200⁰C. Table 1 and Table 2 list the process parameters and chemical composition of powders used in this study for the fabrication of test samples by the LB- PBF and LP-DED process, respectively. The LP-DED bars were fabricated by the RPM Innovations (RPMI), while the LB-PBF bars were produced by SLM Solutions. 5-mm thick test coupons were cut for microstructural analysis. The test coupons were heat treated using an electric furnace with a heating rate of 5⁰C/minute.Detailed descriptions of the multiple-step heat treatment processes used in this study, along with the designations used hereafter in this article to identify each test sample, are listed in Table 3, which are further schematically shown in Figure 1. The heat treatment process could consist of three consecutive steps [15–17].

The heat treatment process could consist of three consecutive steps [15–17]. As Step 1, a stress-relieving (SR) at 1066°C for 1.5 hr followed by air cooling to ambient temperature, and as Step 2, a hot isostatic pressing (HIP) process at 1163°C under 103 MPa isostatic pressure for 3924hr [6,18] followed by argon quench to room temperature, and finally as Step 3, a solution treatment at 1177°C followed by argon quench or water quench [6,8,18] to room temperature.Microstructural coupons were mounted, ground, and polished in the plane perpendicular to the build direction according to ASTM-E3 [19]. A Zeiss Crossbeam 550 scanning electron microscope (SEM) equipped with energy dispersive spectroscopy for (EDS) and Electron backscatter diffraction microstructural analyses. An electron channeling contrast imaging (ECCI) technique [20] was used to obtain the backscattered secondary electron (BSE) micrographs.

An electron channeling contrast imaging (ECCI) technique [20] was used to obtain the backscattered secondary electron (BSE) micrographs. It should be noted that EBSD analysis using a step size of 0.43 µm was conducted for grain size measurement; therefore, grains smaller than 0.43 µm were excluded from the analysis.(EBSD) detectors were usedA Leco LCR500 Rockwell-type hardness tester was used to measure the macro- hardness of the test samples in various heat treatment conditions by a Rockwell B indenter using 100 kgf load according to ASTM E18 standard [21,22]. It should be noted that each mean value reported in this study was calculated over at least 5 hardness readings.Table 1. Process parameters used in this study for fabrication of the LB-PBF and LP- DED Hastelloy-X test samples.Process Power(W)LB-PBF LP-DED200 1070Layer thickness (µm) 30 381Scan speed (mm/sec) 1000 -Travel speed (mm/min) - 1016Hatch distance (µm) 100 -Powder feed rate (g/min)16Table 2.

Schedules of heat treatment conducted on both LB-PBF and LP-DED Hastelloy-X alloy in this study.Results and DiscussionInverse pole figure (IPF) maps obtained from EBSD analysis in the plane perpendicular to the build direction for LB-PBF and LP-DED Hastelloy-X test samples in various heat treatments used in this study are shown in Figure 2. As shown Figure 2, the average grain sizes of the LB-PBF Hastelloy X alloy in various heat treatment conditions, ranging from 46 μm for the non-heat treated (NHT or as-built) sample to 38 μm for the fully heat treated one (Figures 2(a)-(g)), are smaller than those of the LP-DED samples, ranging from 53 μm for No-HIP-No to 65 μm for the fully heat treated conditions (see Figures 2(i)-(n)). The differences in grain size between LB-PBF and LP-DED Hastelloy- X can be attributed to differences in layer height (see Table 1 and Table 2) of LB-PBF and LP-DED processes as well as possible differences in the chemical composition of their powders.

The difference in chemistry may lead to more or less carbides, which are effective grain boundary pinners. It should be noted that the finer grain sizes observed in the LB-PBF alloy may have a slight improving effect on the strength of the alloy as compared with the LP-DED samples, according to the Hall-Petch relationship [23]. Furthermore, it is shown in Figure 3 that the further multiple steps of post heat treatments have a homogenization effect on the grain size distribution of both the LB-PBF and LP- DED samples.It should be noted that the multiple-step heat treatments resulted in a decrease in average grain sizes of the LB-PBF alloy compared to the NHT condition (see Figure 2), while in the case of the LP-DED samples, the subsequent heat treatment has a minimal927effect on the average grain size of the alloy.

This can be further shown in Figure 4, where the BSE micrographs obtained by ECCI technique for both LB-PBF and LP-DED alloys are shown and compared in two magnifications for all the heat treatment conditions investigated in this study. As shown in low magnified BSE images (c) to (g), the subsequent heat treatments above 1066°C (i.e., stress relieving) have successfully removed the prior dendritic microstructure observed in the NHT condition of the LB-PBF Hastelloy X alloy (see in (a)). However, this is not the case for the LP-DED alloy, and it seems that the subsequent heat treatments could only partially dissolve the inter-dendritic regions, as shown in Figures 4(i) to (vii).928Figure 2.

Inverse pole figure (IPF) maps obtained by EBSD analysis on the plane perpendicular to the build direction in (a) to (g) for the LB-PBF Hastelloy-X alloy and in (h) to (n) for the LP-DED samples in various heat treatment conditions investigated in this study.These different effects of heat treatment on the LB-PBF and LP-DED Hastelloy X microstructure could partly be attributed to the differences in the chemical composition, especially the carbon content of powder used for the LB-PBF and LP-DED processes. Specially, the LB-PBF powder batch may have lower carbon content, which can explain the lower fraction of carbides and more recrystallization/growth seen in the heat treated LB-PBF microstructure. Further, as shown in high magnification BSE images, some phases were formed at grain boundaries of both the LB-PBF and LP-DED alloys, as shown in Figures 4(h)–(o) and (viii)–(xiv), respectively.

Further, as shown in high magnification BSE images, some phases were formed at grain boundaries of both the LB-PBF and LP-DED alloys, as shown in Figures 4(h)–(o) and (viii)–(xiv), respectively. For the LB-PBF samples, such phases may be the Cr-rich (M23C6) and Mo-rich (M6C) carbides [2,9,24]. However, for the LP-DED samples, these phases could be identified as the Laves phases, according to Ref [14].Figure 3. Typical comparative column graph for the grain size distributions of LB-PBF and LP-DED Hastelloy-X alloys in various heat treatment conditions investigated in this study in (a)-(b) and (c)-(d), respectively.For parts fabricated using AM technology, it is recommended to conduct stress relief to remove residual stresses [25]. It can be seen from the IPF maps (see Figures 2(b) and (i) and Figure 3) that conducting stress-relieving at 1066⁰C (i.e., 1066-No-No) on both LB-PBF and LP-DED Hastelloy-X partially homogenized the grains.

It has been929reported in the literature that exposure to high temperatures (1066⁰C for 1.5 hours) may cause partial homogenization [25,26]. However, after stress relief, the average grain size is almost unchanged for 1066-No-No LB-PBF and LP-DED Hastelloy-X alloy compared to NHT conditions.Figure 4. Typical BSE micrographs in two magnifications obtained in the plane perpendicular to build direction for the LB-PBF Hastelloy-X in (a) to (o) as well as the930LP-DED Hastelloy-X in (i) to (xiv) in various heat-treated conditions. Red arrows are pointing to the carbides, and blue arrows are pointing to the micro-segregations.The comparative column graph of the macro-hardness measurements obtained by the Rockwell B (HRB) method is presented in Figure 5 for both the LB-PBF and LP-DED alloys in various heat treatment conditions.

As shown, the HRB macro-hardness values of the LB-PBF samples are significantly higher than those of the LP-DED counterparts for all heat treatment conditions that do not include HIP. This could partly be ascribed to the strengthening effect of grain sizes, according to the Hall-Petch relationship [23]. It should be noted that upon the first step low-temperature stress-relieving at 1066°C, the hardness of the LB-PBF samples have not significantly changed, which could be ascribed to the strengthening effect [9] of remnants of the inter-dendritic regions (see Figure 4(b)) as compared with the other higher temperature heat treatments (see Figures 4(b)–(g)).Figure 5. Comparative column graph for the Rockwell B hardness values of LB-PBF and LP-DED Hastelloy-X samples in various heat treatment conditions investigated in this study.Among different heat treatments of LB-PBF Hastelloy-X, NHT and 1066-No-No exhibit the highest hardness (~90 HRB).

The presence of fine grains along with the fine dendritic microstructure might have contributed to the high hardness of LB-PBF Hastelloy- X in NHT condition (see Figure 4(h)). After No-No-1177 and 1066-No-1177, grains of LB- PBF Hastelloy-X have been significantly homogenized, and carbides from inter-dendritic regions have been dissolved (see Figures 4(j) and (l)). The dissolution of dendritic microstructure might result in the decrease in hardness values of No-No-1177 and 1066- No-1177 LB-PBF Hastelloy-X samples (~15%) compared to NHT and 1066-No-No conditions. On the other hand, although No-HIP-No and 1066-HIP-No heat treatments931homogenized the grains, the formation of carbides along grain boundaries (see Figures 4(k) and (m)) increased the indentation force response of LB-PBF Hastelloy-X since carbide strengthen Hastelloy-X. Consequently, No-HIP-No and 1066-HIP-No LB-PBF Hastelloy-X exhibit slightly higher hardness (~6%) compared to 1066-No-No and 1066- No-1177 conditions.

Dissolution of the inter-dendritic region, as well as partial dissolution of inter-dendritic carbides in No-No-1177 and 1066-No-1177 LP-DED Hastelloy-X (see Figures 4(x) and (xii)), resulted in the lowest hardness values among all other conditions. After No-HIP-No and 1066-HIP-No post-processing conditions, the phases in the prior inter-dendritic regions were well dissolved which may have improved the degree of solid solution and increased their strength (~8% increase in HRB). These resulted in the highest hardness (~81 HRB) of LP-DED Hastelloy-X in No-HIP-No and 1066-HIP-No among all the heat treated LP-DED Hastelloy-X. The three HIPed conditions did not yield similar hardness values for LP-DED samples.

The three HIPed conditions did not yield similar hardness values for LP-DED samples. It is also interesting to note that both LB-PBF and LP-DED materials had quite comparable hardness values in HIPed conditions.ConclusionsIn this study, the effect of multiple-step heat treatment on microstructure and mechanical properties of the Hastelloy X fabricated using two different AM processes, including LB-PBF and LP-DED are investigated and compared. The test samples underwent multiple combinations of heat treatments such as stress-relieving (SR) at 1066°C for 1.5 hr, followed by hot isostatic pressing (HIP) at 1163°C at103 MPa for 3 hr, and a solution heat treatment at 1177°C for 3 hr. The microstructure as the result of each single- and/or multi-stage heat treatment was characterized. Further, the macro-hardness of the test samples was measured using the Rockwell B (HRB) method.

The microstructure as the result of each single- and/or multi-stage heat treatment was characterized. Further, the macro-hardness of the test samples was measured using the Rockwell B (HRB) method. Below is a brief summary of experimental observations from this study:The EBSD analysis reveals that the subsequent heat treatments have a homogenization effect on the grain structure of both the LB-PBF and LP-DED materials; however, the latter has shown larger grain sizes in various hat treatments as compared with the former, which could partly be ascribed higher layer height used in the RPM Innovation machine for fabrication of the LP-DED samples as well as possible differences in the chemical composition of powders used in LB-PBF and LP-DED processes.The BSE micrographs reveal that the prior dendritic microstructure of the as- deposited LP-DED samples are hardly dissolved upon subsequent heat932treatments.

However, for the LB-PBF samples, the prior inter-dendritic regions are almost completely removed upon the subsequent heat treatments except for the single-step stress relieving at 1066°C (1066-No-No).The HRB hardness measurements show that the LP-DED samples show lower HRB hardness values than those of the LB-PBF counterparts, which could be attributed to grain size strengthening.Upon subsequent heat treatments without HIP, both materials have shown a softening effect. However, the hardness of the LB-PBF samples was significantly higher in NHT and stress-relieved (SR at 1066°C) conditions, as compared with that of solution treated samples at 1177°C, which could be attributed to the strengthening effect of the much finer prior dendritic microstructure. AcknowledgmentThis work is supported by the National Aeronautics and Space Administration through Grant# 80MSFC19C0010. The authors would like to thank SLM Solutions and RPMI for fabrication support of samples in this study.

The authors would like to thank SLM Solutions and RPMI for fabrication support of samples in this study. This paper describes objective technical results and analysis. Any subjective views or opinions that might be expressed in the paper do not necessarily represent the views of the National Aeronautics and Space Administration (NASA) or the United States Government.References[1] Haynes Internatioal Website, Hastelloy X: Tensile Properties. [2] M. Aghaie-Khafri, N. Golarzi, Forming behavior and workability of Hastelloy Xsuperalloy during hot deformation, Mater. Sci. Eng. A. 486 (2008) 641–647.[3] S. Bremen, W. Meiners, A. Diatlov, Selective Laser Melting, Laser Tech. J. 9(2012) 33–38.[4] M. Molitch-Hou, Overview of additive manufacturing process, Addit. Manuf. (2018)1–38.[5] P.D. Nezhadfar, R. Shrestha, N. Phan, N. Shamsaei, Fatigue behavior ofadditively manufactured 17-4 PH stainless steel: Synergistic effects of surface roughness and heat treatment, Int. J.

Rometsch, M. Heilmaier, X. Wu, Influence of post heat treatments on anisotropy of mechanical behaviour and microstructure of Hastelloy-X parts produced by selective laser melting, Mater. Sci. Eng. A. 667 (2016) 42–53.[13] O. Sanchez-Mata, J.A. Muñiz-Lerma, X. Wang, S.E. Atabay, M. Attarian Shandiz, M. Brochu, Microstructure and mechanical properties at room and elevated temperature of crack-free Hastelloy X fabricated by laser powder bed fusion, Mater. Sci. Eng. A. 780 (2020) 139177.[14] W. Zhang, Y. Zheng, F. Liu, D. Wang, F. Liu, C. Huang, Q. Li, X. Lin, W. Huang, Effect of solution temperature on the microstructure and mechanical properties of Hastelloy X superalloy fabricated by laser directed energy deposition, Mater. Sci. Eng. A. 820 (2021) 141537.[15] ASTM F3301, Standard for Additive Manufacturing – Post Processing Methods – Standard Specification for Thermal Post-Processing Metal Parts Made Via Powder Bed Fusion, ASTM Stand. (2018) 3.[16] H. V Atkinson, S.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAdditively manufactured Haynes 230 by laser powder directed energy deposition (LP-DED): effect of heat treatment on microstructure and tensile propertiesMuztahid Muhammad1,2, Reza Ghiaasiaan1,2, Paul R. Gradl3, Shuai Shao1,2, Nima Shamsaei1,2*1National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849, USA 2Department of Mechanical Engineering, Auburn University, Auburn, AL 36849, USA 3NASA Marshall Space Flight Center, Propulsion Department, Huntsville, AL 35812, USACorresponding author: shamsaei@auburn.edu Phone: (334) 844-4839AbstractThe microstructure and tensile mechanical properties of Haynes 230 fabricated through laser powder directed energy deposition (LP-DED) were investigated, varying temperature heat treatments between 900°C to 1177°C following deposition.

Scanning electron microscopy (SEM) was employed for microstructural analysis, whilst tensile testing was utilized to evaluate the room temperature mechanical properties of the alloy. In an as- deposited state, the initial microstructure consisted of cellular γ and M6C/M23C6 carbides. The cellular regions seem to be fully dissolved upon solutionizing at 1177°C for 3 hours. Following post-deposition heat-treatments, the carbides were observed to precipitate and grow along the grain boundaries as well as in the interior of grains. Solutionizing at 1177°C for 3 hours following stress-relieving yielded better ductility and had an insignificant effect on the strength.Keywords: Laser powder directed energy deposition, Haynes 230, microstructure, tensile behavior.IntroductionHaynes 230 is a solid solution and carbide strengthening nickel-base superalloy with high-temperature mechanical strength and corrosion resistance.

This alloy is used in many aerospace applications, therefore a potential candidate for applications such as hot sections for gas turbines, heat exchangers of nuclear reactors, etc.[1]. However, machining of Haynes 230 alloy is a challenge, especially for components with complex geometries. Additive manufacturing (AM), a layer-by-layer manufacturing technology, has the capability to fabricate complex parts in near-net-shape conditions with none to minimal machining requirements.

Additive manufacturing (AM), a layer-by-layer manufacturing technology, has the capability to fabricate complex parts in near-net-shape conditions with none to minimal machining requirements. Among different AM processes, laser powder directed energy deposition (LP- DED) is a well-known AM process where injected powder particles are melted together by a laser heat source and are deposited into the melt pool [2].Although AM offers numerous benefits, there are challenges associated due to AM935parts’ unique thermal histories, such as rapid melting and solidification, leading to different microstructure compared to the conventionally fabricated parts as well as induced residual stress and volumetric defects [3]. Consequently, the mechanical properties of AM parts may differ from their conventionally manufactured counterparts.

Consequently, the mechanical properties of AM parts may differ from their conventionally manufactured counterparts. Hence, the microstructure and mechanical properties of AM parts need to be studied before the implementation of heat treatments designed for wrought materials.In general, as-fabricated AM Ni-base superalloys consist of cellular inter-dendritic structures that affect mechanical properties compared to conventionally fabricated counterparts. Solid solution and carbide strengthened Ni-base superalloys fabricated using both wrought [4] and AM [5] processes are known for their superior ductility.

Solid solution and carbide strengthened Ni-base superalloys fabricated using both wrought [4] and AM [5] processes are known for their superior ductility. Since solid solution strengthened Ni-base superalloys such as Haynes 230 do not possess detrimental Nb in their chemical composition like precipitation strengthened Ni-base superalloys, the Haynes 230 has the potential to improve the ductility by post-process homogenization (or partial removal) of the initial micro-segregation and inter-dendritic structure induced during the prior AM fabrication process with a minimal negative effect on strengths. Although Haynes 230 alloy has been studied in conventionally processed [1,4,6–8] and AM conditions [9,10], there is a lack of knowledge for the effect of multiple-step heat treatment processes on subsequent mechanical properties of AM Haynes 230 alloy. In this study, the effect of different heat treatments on the microstructure and tensile properties of LP-DED Haynes 230 is investigated.

In this study, the effect of different heat treatments on the microstructure and tensile properties of LP-DED Haynes 230 is investigated. Tensile behaviors of LP-DED Haynes 230 at various heat treatment conditions are correlated with their corresponding microstructure.Experimental ProcedureThe Haynes 230 test specimens were fabricated using the RPM Innovations (RPMI) 557 LP-DED system using the process parameters tabulated in Table 1. Powders used to fabricate the Haynes 230 specimens were supplied by Praxair Surface Technologies, Inc., and the chemical composition of powders is listed in Table 2. The AM test specimens were machined to a final geometry of tensile specimens according to ASTM E8 standard [11], as shown in Figure 1. Prior to the final machining, the tensile test specimens were heat treated using a ‘Sentro tech high-temperature tube furnace’ with a heating rate of 5⁰C/minute.

Prior to the final machining, the tensile test specimens were heat treated using a ‘Sentro tech high-temperature tube furnace’ with a heating rate of 5⁰C/minute. To investigate the effect of prior stress-relieving (SR), the specimens were stress-relieved (SR) at two different temperatures, i.e., 900⁰C and 1066⁰C for 1.5 hours, followed by solutionizing at 1177⁰C for 3 hours, as shown in Figure 2 and Table 3. The suggested range of solution treatment for wrought Haynes 230 alloy is from 1177⁰C to 1246⁰C [4,12–14].Microstructural analyses were performed on small coupons cut from heat treated test specimens, in the plane perpendicular to the build direction. The microstructural coupons were later mounted and ground, and polished according to ASTM-E3 [15]. A Zeiss Crossbeam 550 scanning electron microscope (SEM) equipped with energy dispersive spectroscopy (EDS) and Electron backscatter diffraction (EBSD) detectors were used for further microstructural analysis.

Backscattered secondary electron (BSE) micrographs were obtained using the electron channeling contrast imaging (ECCI) technique [16].Quasi-static monotonic tensile testing was conducted utilizing an MTS landmark936servo-hydraulic machine with a load cell of 100 KN. Three tensile specimens were tested for each heat treated condition. An extensometer was used for accurate yield strength measurement during tensile testing. However, for the limited displacement range of the extensometer, the tensile test was conducted in two stages. The extensometer was attached to the surface of the specimen only up to the first 0.05 mm/mm strain of the tensile test. After removal of the extensometer, the remaining test was conducted in displacement-controlled mode until fracture.Table 1. Process parameters used in this study for fabrication of the LP-DED Haynes 230 specimens.Power (W)1070Layer height (µm)381Travel speed (mm/min) 1016Powder feed rate (g/min)19.10Table 2.

Tensile specimen geometry of LP-DED Haynes 230 (All dimensions are in mm).Figure 2. Diagram presenting heat treatments conducted on LP-DED Haynes 230 in this study.Table 3. Heat treatments (with designations) conducted on LP-DED Haynes 230 in this study.Heat treatmentStep 1 (Stress-relieving)Temp- erature (⁰C)Time (hr)Coolin gStep 2 (Soltutionizing)Temp- erature (⁰C)Time (hr)Coolin gDesignatio nAs-builtNoneNoneNoneNon-heat treated (NHT)Stress- relievedStress- releived + Solutionize d900⁰C1066⁰C900⁰C1066⁰C2Air (AC)None1177⁰CNone None3Water (WC)900-None1066-None900-11771066-1177938Results and DiscussionInverse pole figure (IPF) maps obtained from EBSD analysis for all heat treatment conditions are presented in Figure 3 along with their grain size distribution in the plane perpendicular to the build direction. As shown, typical fine equiaxed grains are observed in the non-heat treated (NHT) condition.

As shown, typical fine equiaxed grains are observed in the non-heat treated (NHT) condition. Furthermore, some ultra-fine grain regions are also discernible in the EBSD map, which could be ascribed to the inhomogeneous formation of grain boundary carbides pinning the growth of the ultra-fine grains. These carbides are indeed effective in resisting grain growth as there are not much recrystallization and grain growth observed even after further heat treatments, as shown in Figure 3. Note that it is expected that the NHT condition has more small grains which gradually grow during heat treatment even though it is indicated otherwise in Figure 3(f). These grains may be too small for the IPF mapping to capture at the magnification used (i.e., the probe points have a much higher probability to be near grain boundaries), which led to a higher fraction of zero solutions during EBSD analysis.939Figure 3.

Inverse pole figure (IPF) maps of LP-DED Haynes 230 in the plane perpendicular to the build direction for different heat treated conditions: (a) NHT, (b) 900- None, (c) 1066-None, (d) 900-1177, and (e) 1066-1177; (f) Grain size distribution of LP- DED Haynes 230 for different heat treated conditionsThe higher resolution, BSE micrographs for all heat treatments investigated are presented in Figure 4. It can be seen that the NHT DED Haynes 230 consists of cellular dendritic microstructure with severe micro-segregation in the ɣ matrix, as shown in Figures 4(a)-(b), which is consistent with the reported data in the literature [9]. Typically, cellular inter- dendritic network in the AM materials can increase the strengths at the cost of the ductility. Formation of some filmy carbides along the inter-dendritic regions is also observed at the940NHT LP-DED Haynes 230 (indicated by red arrows in Figure 4(b)).

Formation of some filmy carbides along the inter-dendritic regions is also observed at the940NHT LP-DED Haynes 230 (indicated by red arrows in Figure 4(b)). Chemical composition observed by EDS analysis in Haynes 230 indicates the presence of Mo-rich and Cr-rich carbides (as shown in Figure 5), which may be Mo-rich (M6C) and Cr-rich (M23C6) carbides typical for Haynes 230 and is consistent reports in the literature [8,9]. Carbides also can contribute to the high-temperature strengthening mechanism in the Haynes 230 by resisting the grain boundary sliding at high temperatures [17].Figure 4. Typical BSE micrographs obtained in the plane perpendicular to build direction for various heat treated conditions of the LP-DED Haynes 230: (a)-(b) NHT, (c)-(d) 900-None, (e)-(f) 1066-None, (g)-(h) 900-1177, and (i)-(j) 1066-1177. Red and941blue arrows point to the carbides formed at grain boundaries and grain interior regions, respectively.Figure 5.

Red and941blue arrows point to the carbides formed at grain boundaries and grain interior regions, respectively.Figure 5. Typical elemental maps and the overall chemical content from EDS for the 1066- 1177 solution treated LP-DED Haynes 230 specimen in the plane perpendicular to the build direction.Upon the low-temperature 1-step heat treatment (or SR) at 900°C (900-None), the micro-segregations at the inter-dendritic regions are dissolved and permitted the formation of more filmy carbides at these locations, as shown in Figures 4(d). For the wrought Haynes 230 [8], large carbides (i.e., M6C) start dissolving at temperature 600⁰C, and smaller M23C6 carbides start forming in grain interior regions.

For the wrought Haynes 230 [8], large carbides (i.e., M6C) start dissolving at temperature 600⁰C, and smaller M23C6 carbides start forming in grain interior regions. These small carbides forming with the grain interior regions could play as a minor strengthening source for the alloy [17].On the other hand, upon the high-temperature 1-step heat treatment (or SR) at 1066°C (1066-None), the inter-dendritic regions are more significantly dissolved as compared to the low-temperature SR at 900°C (900-None), as shown in Figures 4(e)-(f). It is also notable that in 1066-None specimens, the grain boundary and inter-dendritic carbides942have become more globular and the number density of the carbides in the grain interior regions are increased (carbides at grain interior regions are indicated by blue arrows in Figure 4(f)), which could be ascribed to the prior inter-dendritic carbides that were dissolved after heat treatment or diffusion of elements at the grain interior regions.

Furthermore, it seems that the 2-step heat treatment (or solution treatment) at 1177°C could completely remove the inter-dendritic regions, as shown in Figures 4(g)-(j). It should be noted that, upon 2-step heat treatment (or solution treatment) at 1177°C, (900-1177 and 1066-1177), the grain boundary carbides tend to become discrete and spheroidize (See Figures 4(h) and (j)).Engineering stress-engineering strain and engineering stress-displacement curves of LP-DED Haynes 230 in various heat treatment conditions investigated in this study are shown in Figures 6(a) and (b), respectively.

In addition, Figure 7 presents the comparative column charts for uniaxial tensile properties, including yield strength (Sy), ultimate tensile strength (Su), and percent elongation (%EL) of LP-DED Haynes 230 investigated in this study with and those of laser powder bed fusion (L-PBF) [10] and wrought [4] counterparts found in the literature.As shown in Figure 6 and Figure 7, upon the 1-step heat treatment (or SR), the elongation has not significantly changed as compared with the NHT condition (~36-38%). However, the higher temperature 1-step heat treatment (or SR) at 1066°C (1066-None) has lowered the yield strength (~353 MPa) as compared with the NHT conditions (~433 MPa), whereas the SR at 900°C (900-None) has not significantly affected the yield strength (~431 MPa). This could be ascribed to the effect of 1-step heat treatment on microstructure.

This could be ascribed to the effect of 1-step heat treatment on microstructure. As shown in Figures 4(c)-(d), the lower temperature 1-step heat treatment (900-None) has partially removed the micro-segregation within the prior inter-dendritic regions and formed more carbides in the grain interior. These carbides can lead to more pronounced dislocation multiplication and greater strain hardening (compare the curves for NHT and 900-None in Figure 6(b)). On the other hand, the higher temperature 1-step heat treatment (1066-None) not only succeeded in complete dissolution of prior micro-segregation but also could partially remove the prior inter-dendritic regions. In addition, some carbides formed within the grain interior regions.

In addition, some carbides formed within the grain interior regions. Therefore, the lower yield strength after the high-temperature 1-step heat treatment (1066-None) could be attributed to the partial removal of the inter-dendritic structures, as shown in Figures 4(c)-(d), which could play as impediments for the dislocations’ motion during deformation [5].Furthermore, upon the 2-step heat treatment at 1177°C, the uniaxial tensile properties were comparable with a slight increase in ductility. This could be attributed to the homogenization effect of high-temperature 2-step heat treatment on the microstructure where the carbides formed at grain boundary and grain interior regions have appeared in a more circular morphology, and almost all the prior inter-dendritic regions are removed.943Figure 6. Tensile behavior of LP-DED Haynes 230 alloy: (a) engineering stress-engineering strain curves, and (b) engineering stress-displacement curves.944Figure 7.

Tensile behavior of LP-DED Haynes 230 alloy: (a) engineering stress-engineering strain curves, and (b) engineering stress-displacement curves.944Figure 7. Column chart presenting tensile properties of LP-DED Haynes 230 alloy investigated in this study as well as L-PBF [10] and wrought [4] Haynes 230 alloys from literature.Summary and ConclusionsIn this study, the microstructure and tensile behavior of LP-DED Haynes 230 alloy after various multiple steps heat treatments were investigated. The tensile behavior of each heat treatment was characterized and discussed in relation to their microstructure. The effect of hot isostatic pressing (HIP) in the microstructure and mechanical properties of AM Haynes 230 alloy will be explored in the future.

The effect of hot isostatic pressing (HIP) in the microstructure and mechanical properties of AM Haynes 230 alloy will be explored in the future. The following conclusions can be drawn from this study:The microstructural results revealed that the low-temperature 1-step heat treatment (SR) at 900°C has partially dissolved the prior micro-segregation with insignificant effect on the inter-dendritic regions while the carbides formed at grain boundaries become more discrete and continuous at grain boundary areas.

On the other hand, the higher temperature 1-step heat treatment (SR) at 1066°C has not only removed almost all the prior micro-segregation but also partially dissolve the prior inter-dendritic structure.Upon the high-temperature 2-step heat treatment (or solution treatment) at 1177°C, the microstructure has become more homogenized with the partial dissolution of the inter- dendritic regions as well as the formation of discrete carbides at grain interior and boundary areas.The tensile results of the low-temperature 1-step heat treatment (SR) at 900°C showed comparable properties with non-heat treated (NHT) conditions which could945be attributed to the minimal effect of heat treatment on prior inter-dendritic regions, which could play as impediments for dislocation motions.

On the other hand, the high- temperature 1-step heat treatment (SR) at 1066°C has lowered yield strength due to partial removal of inter-dendritic regions.The 2-step heat treatment at 1177°C, the tensile properties are more comparable regardless of prior 1-step heat treatment, which could be attributed to the homogenization effect of the 2-step heat treatment at 1177°C on microstructure where the inter-dendritic regions are partially removed while the carbides are formed at the grain boundary and interior regions in a more circular shape. AcknowledgmentThis material is partially supported by the National Aeronautics and Space Administration NASA under Cooperative Agreement No. 80MSFC19C0010. This paper describes objective technical results and analysis. Any subjective views or opinions that might be expressed in the paper do not necessarily represent the views of the National Aeronautics and Space Administration (NASA) or the United States Government.References[1] C.J.

HESIVE BONDING OF SHEET FOR LAMINATED METAL TOOLINGGene Zak and Wendy Xu WangCentre for Automotive Materials and Manufacturing Department of Mechanical Engineering, Queen's University, Kingston, ON, Canada K7L 3N6AbstractThere exists a significant body of work on metal laminate tooling built by the “cut-stack-bond” approach; however, automation with this method is difficult. Building laminations by "stack-bond-cut" sequence, on the other hand, is more amenable to automation. Two main challenges of "stack-bond-cut" sequence are blind contour cutting and bonding of the sheet. In this study, we investigate the hot-roller method of thermoplastic adhesive bonding for the metal laminations. Metal sheet, having thermal characteristics significantly different from paper, poses its own specific problems. During the bonding process, in order to achieve good bond strength, appropriate heat and pressure must be applied. As the stack builds up, thermal and mechanical properties change.

During the bonding process, in order to achieve good bond strength, appropriate heat and pressure must be applied. As the stack builds up, thermal and mechanical properties change. This inconsistency of process conditions can potentially lead to part warpage, unless carefully controlled. Temperature measurements with a thermocouple embedded into lamination stack showed the effect of bonding process parameters on the laminate temperature.IntroductionUsing laminations (i.e., material in sheet form) for layered manufacturing (LM) avoids phase transformations with attendant shrinkage (as in powder sintering) and allows easy process scale-up. Adams et al. [1] produce working tools from steel laminations 1-3 mm thick, but have to resort to a finish milling step to achieve acceptable surface quality. Himmer et al. [2] propose to fabricate injection moulds by (1) joining flux-coated 2.5-mm-gauge aluminum through heat treatment near the flux melting temperature and (2) finish milling.

Himmer et al. [2] propose to fabricate injection moulds by (1) joining flux-coated 2.5-mm-gauge aluminum through heat treatment near the flux melting temperature and (2) finish milling. Walczyk and Dolar [3] present results of tests on several epoxy and alumina-based adhesives used to bond aluminum laminations.The “cut-stack-bond” approach (Figure 1a) adopted by above authors involves (1) cutting out the desired part’s profile at each lamination; (2) assembly (stacking) of the profiles, and (3) bonding (mechanically with fasteners, adhesively, or by welding). The primary drawback of this approach is the complexity of handling the cut-out profiles between the cutting and stacking steps, requiring indexing, transporting, and alignment. This complexity makes automation of the process highly impractical. An easier-to-automate alternative is a “stack-bond-cut” sequence (Figure 1b).

This complexity makes automation of the process highly impractical. An easier-to-automate alternative is a “stack-bond-cut” sequence (Figure 1b). This method has been applied in Laminated Object Manufacturing (LOM), but the sheet materials have been limited to paper, polymers, and ceramic composites [4]. Replacing soft materials with metals is a challenge, however, due to difficulties with “blind” cutting of metals (i.e., cutting only to the depth of the sheet’s gauge) [5]. This paper presents results of experimental investigation into thermal mechanism of adhesive bonding of metal laminations.502CUTSTACK(a)BONDSTACKBOND (b)Figure 1. Two possible sequences for producing laminated objects: (a) cut-stack-bond and (b) stack- bond-cut.Bonding of Metal LaminatesThere are several alternatives for metal sheet-to-sheet bonding. First, heat-activated polymeric adhesive (used in paper-based LOM process) can be also applied to metals.

First, heat-activated polymeric adhesive (used in paper-based LOM process) can be also applied to metals. This kind of bonding is the easiest to implement in an automated process since the adhesive can be pre-applied to the sheet and subsequently quickly activated by heat and pressure. Other methods, such as welding, soldering and brazing may offer higher strength and high-temperature resistance, but they require expensive equipment and complicated processes. Therefore, polymeric adhesive bonding was selected for our initial trials.In order to bond two solid sheets, a liquid adhesive should be spread and solidified between the two surfaces. The reason this results in a strong bond is that the liquid conforms to the surface shapes of the two solids being joined. This permits close contact between the surface of the liquid (adhesive) and the surface of the solids (adherends).

This permits close contact between the surface of the liquid (adhesive) and the surface of the solids (adherends). Such close contact is necessary since the forces which act to hold the matter together are the same ones which to a large extent are responsible for the strength of the adhesive joints.The process of adhesive application usually consists of spreading a thin layer of liquid adhesive between the surfaces to be attached and pressing these surfaces together. A joint is formed when the viscosity of the liquid layer is increased by such mechanisms as solvent permeation or evaporation, polymerization, and cooling until solidification. The requirement for optimum adhesion is good wetting of metal sheet by adhesive. In order to achieve good wetting, appropriate heat and pressure should be applied since heat lowers the viscosity of adhesive and pressure helps to spread the adhesive onto the surfaces.

In order to achieve good wetting, appropriate heat and pressure should be applied since heat lowers the viscosity of adhesive and pressure helps to spread the adhesive onto the surfaces. Insufficient heat/pressure may cause weak interlaminar bonds or even delamination. Excessive heat/pressure may cause the material degradation. Therefore, it is important to control accurately the heat and pressure in bonding process.Besides bond strength, other factors, such as cure speed, cure method and suitability for use in automated production, will all play a role in determining the best adhesive for a specific application. First, because thin-gauge (0.1-0.5mm) aluminum or steel sheet will be used as the laminate, the adhesive layer must be kept to a minimum thickness. Second, high bond strength and good temperature resistance are important because of the intended functional tooling applications.Two basic types of adhesives were considered: solvent-based and film.

Solvent-based adhesives are spread evenly along the material to be bonded before the solvent or water is evaporated to create a bond. The adhesive can then be activated using heat and pressure. Film adhesives facilitate manual503CUTapplication since they result in uniform, controlled bond-line thickness. Film adhesives were used in the tests reported herein.Literature SurveyA heated roller is used in paper-based LOM to apply the heat and pressure to the laminate. For our first prototype, we have adopted a similar system. There have not been to the authors’ knowledge reports of thermal bonding investigations for thin metal laminates. Following is a brief overview of such work for the non-metallic laminates.Park et al. [6] determine the optimum range for the heater temperature and speed settings using a commercial LOM machine by conducting tensile tests on bonded paper specimens.

[6] determine the optimum range for the heater temperature and speed settings using a commercial LOM machine by conducting tensile tests on bonded paper specimens. Chen [7] measures the heat transferred across the paper sheet by placing a thermocouple underneath it. Results are summarized in paper bonding curves. Each curve represents temperature rise in the adhesive as a function of roller speed at a fixed roller temperature. Curves show that the temperature rise decreases with increasing speed and decreasing roller temperature. It was also noted that the effect of roller pressure is not very pronounced.Pak and Nisnevich [8] develop a mathematical model of the relationship between temperature at the arbitrary depth of the part and the LOM-process parameters such as temperature, speed of the heater, and pressure exerted by the heated roller. Flach et al. [9] also develop a mathematical model that describes the heat transfer during building of ceramic LOM parts.

Flach et al. [9] also develop a mathematical model that describes the heat transfer during building of ceramic LOM parts. The thermal behavior of a part is determined by heat transfer to the part from the roller, heat conduction within the part itself, heat loss from the part to the metal base plate, and heat loss to the surroundings. The appropriate differential equation and boundary conditions were set up for a rectangular geometry part and then, together with appropriate initial conditions, material properties, and heat transfer coefficients, numerically solved to reveal the temperature distribution within the part. By tuning one of the model parameters (roller-to-part heat transfer coefficient), good agreement was obtained between predicted temperatures within a part and the actual temperature as measured by embedded thermocouples. In Flach et al. [10] the above model is used to investigate the effects of varying certain parameters during a number of simulated LOM builds.

In Flach et al. [10] the above model is used to investigate the effects of varying certain parameters during a number of simulated LOM builds. It is found that chamber air temperature and base-plate temperature combination are the most effective in controlling of the overall part temperature. The roller temperature proved to be the most effective in determining the temperature of the surface layers of the part.Sonmez and Hahn [11] use Finite Element method to analyze the stress and heat transfer distribution in paper-based LOM. Stress analysis results provide the heat transfer model with an estimate of the contact area between the roller and the laminate. Effect of roller radius, laminate built-up thickness, and roller speed on the thermal time history is examined.

Effect of roller radius, laminate built-up thickness, and roller speed on the thermal time history is examined. They find that thickness of the built-up laminate impacts the stress distribution, larger diamter results in greater contact area, and that roller speed must be below certain value for successful bonding.Bonding Temperature Measurement Experiments Objectives: These experiments aimed to measure the temperature profile in the steel laminate while a heated roller is passing by at different rolling speeds, pressures and temperatures. This information would help in setting appropriate process parameters for metal laminate bonding.504Equipment: Figure 2 shows the heated roller used in the experiments to apply the pressure and heat in order to bond the metal laminations. The roller is carried by two supports which are mounted onto a belt-driven linear stage.

The roller is carried by two supports which are mounted onto a belt-driven linear stage. One axis is driven by a stepper motor via Gemini GV motion controller while the other axis is mechanically linked to it by a drive shaft to assure synchronous movement. The pressure is controlled by adjustment of two springs which transfer the horizontal spring force to a vertical pressure via two L-shaped pivoting arms. By varying the spring compression from 0 to 40 mm, the roller downward force can be adjusted from 11 to 65 kg. A 650-Watt rod-shaped heating element located along the centre axis of the roller is PID controlled to within ±0.1°C using feedback provided by a built-in thermocouple sensor.Heated RollerCompression SpringSteel LaminateL-shaped PivotZ PlatformLinear Motion StageRoller Support SlideFigure 2. Heated roller and steel laminate.Laminate Material: The laminate comprised 0.12-mm plain-steel shim stock and Scotch 467MP film adhesive from 3M.

Heated roller and steel laminate.Laminate Material: The laminate comprised 0.12-mm plain-steel shim stock and Scotch 467MP film adhesive from 3M. The 467MP is made from 200MP high-performance acrylic adhesive (50 microns) and polycoated Kraft paper liner. Temperature measurements were made during the building of a laminated injection mould insert (Figure 3a). A half-inch thick aluminum alloy block served as the base (Figure 3b). In the middle of the block, a square slot was cut through to assist in extraction of the laminate at the end of the build. The first layer of steel sheet was fixed to the block by gluing the four edges to four sides of the slot. The square block cut out from the original shape was then reinserted within the opening to provide support. When building was completed, the cut-out block was removed and the opening used to help in removal of the laminate form the base.(a)(b)Figure 3.

When building was completed, the cut-out block was removed and the opening used to help in removal of the laminate form the base.(a)(b)Figure 3. (a) CAD model of laminated injection mould insert (16.5 H mm, 91 layers); (b) support base.505Test Procedure: Two wires of a J-type precision fine-wire (0.12 mm diameter) Teflon-insulated thermocouple were spot welded to the underside of the steel sheet individually so that they are nearly touching (Figure 4a). The wire leads were glued to the sheet to prevent them from moving during bonding. A channel (1.07H×1.5W mm) built within the laminate in order to accommodate the wires (Figure 4b) was modeled in CAD as part of the mould and built automatically. For each laminate which included the channel outline, the cut-out part of the sheet was manually removed. The data acquisition system comprised a NI-DAQ board 6035E, a dedicated temperature signal-processing module 5B37, and a LabVIEW v5.1 software.

The data acquisition system comprised a NI-DAQ board 6035E, a dedicated temperature signal-processing module 5B37, and a LabVIEW v5.1 software. Measurements were collected at 50 samples per second.(a)(b)Figure 4. (a) thermocouple attached to the steel sheet; (b) laminate with thermocouple channel.The process parameters varied included the roller load (L = 12, 25, 40, 51 kg), temperature (Ts = 120, 170, 220, 270, 320°C), and speed (V = 5, 10, 20, 40 mm/s). Thermocouples were embedded in 7th, 45th, and 87th layers. Only results for the 45th layer will be presented. Temperature was measured during the lamination of the sheet to which the thermocouple was attached as well as during the lamination of the subsequent sheets. To attach the sheet with the thermocouple, the film adhesive was spread manually on the previously built stack, its liner was peeled off, and the sheet with the thermocouple was placed so that the wires on its underside fit within the cutout slot.

The roller temperature was allowed to stabilize for 20 minutes after the desired setting was indicated by the heater’s controller. Then the roller was moved across the laminate surface while the data were collected.Results and DiscussionEffect of Roller Load: The roller temperature and speed were kept at a fixed value, while the load was adjusted to observe the effect on the laminate temperature (Figure 5). The sharp peak corresponds to the roller passing directly over the thermocouple. Generally, higher pressure is expected to increase the heat transfer coefficient by increasing the contact surface area. We can observe this effect when the load is increased from lowest setting of 12 kg to the next higher setting of 25 kg. Further increase does not result in proportionate temperature rise. Therefore, only certain minimum pressure is needed to achieve good thermal contact.

Further increase does not result in proportionate temperature rise. Therefore, only certain minimum pressure is needed to achieve good thermal contact. Load was set to 40 kg henceforth.Effect of Roller Temperature and Speed: Good bonding requires certain heat, pressure and dwell time. Thus, the results are presented here as the average temperatures recorded between the time of reaching the peak temperature and five seconds later. As expected, increasing roller temperature setting, leads to proportionate increase in laminate temperature (Figure 6). The proportionality constant increases as the rolling speed decreases. The same information can be also plotted as a function of the roller speed (Figure 7) to obtain the “bonding curves” [7].506200180)160C °(140Load = 12 kge r u t a r e p m e T12010080Load = 25 kgLoad = 40 kgLoad = 51 kg604020051015 Time (s)202530Figure 5.

Effect of roller load on temperature (Layer = 45, Ts = 220 °C, V = 10mm/s).r o f200.p m e Te g a r e v A e c a f r u S e t a n m a Li)C ° (k a e P r e t f as d n o c e S 5180160140120100806040V = 5 mm/sV = 10 mm/sV = 20 mm/sV = 40 mm/s20100150200250300350400Roller's Set Temperature (°C )Figure 6. Effect of roller temperature on the time-averaged steel laminate temperature.r o f200.p m e T)C ° (180160e g a r e v A e c a f r u S e t a n m a Lik a e P r e t f as d n o c e S 5140120100806040Ts = 320 °C Ts = 270 °C Ts = 220 °C Ts = 170 °C Ts = 120 °C2005101520253035404550Rolling Speed (mm/s)Figure 7. Effect of roller speed on the time-averaged steel laminate temperature.507Temperature during the Lamination of Subsequent Sheets: Temperature measurements were continued to be performed as the laminated object was built up above the 45th layer.

In Figures 8 and 9, 1st layer refers to observations during the lamination of 45th layer (to which the thermocouple was attached), 2nd layer value to those during the lamination of the 46th layer and so on. Peak temperature still rises above 100°C for the next two layers after the 45th and then gradually declines. Thus, further significant heating of the adhesive occurs as new sheets are added which is similar to the results obtained in [10]. This implies that the adhesive in upper layers of the part will not have the same degree of heat activation as in the lower layers.2001st layer1802nd layer3rd layer1604th layer)C °(1405th layer6th layere r u t a r e p m e T120100807th layer9th layer11th layer14th layer18th layer22nd layer6032nd layer40200102030405060Time (s)Figure 8.

Temperature recorded during lamination of subsequent sheets (Ts=320°C, V=5 mm/s).200)180C ° (160e r u t a r e p m e T14012010080k a e P6040200246810 12 14 16 18 20 22 24 26 28 30 32Built-up LayersFigure 9. Peak temperatures observed during the lamination of subsequent layers.ConclusionsHigher roller temperature, lower speed and higher roller pressure all resulted in increased heat transfer rate and therefore in higher surface temperature and longer duration of high temperature in the adhesive. As expected, the roller’s temperature and speed are the most important factors. Once sufficient roller load was applied, further increases did not bring improved heat transfer to the laminate. The measurements also showed that the heat dissipates very quickly within the metal laminate. An enclosing508chamber kept at elevated temperature would help to maintain the laminate at a higher temperature for a longer time.

An enclosing508chamber kept at elevated temperature would help to maintain the laminate at a higher temperature for a longer time. Further significant heating of the adhesive was found to occur for several subsequent laminations, which implies that process parameters may have to be adjusted for the upper part layer to assure the adhesive receives similar heat input as on the lower layers.AcknowledgementsWe gratefully acknowledge the financial support of Centre for Automotive Materials and Manufacturing, Ontario Innovation Trust, and Canada Foundation for Innovation. We also wish to express our thanks for assistance with thermocouple mounting to Mr. Claude Quenneville of Alcan Research Centre.References[1] C. Adams, B. Bryden, D. Wimpenny, “Rapid laminated tooling – making the technology work,” 8th European Conf. Rapid Prototyping and Manufacturing, Nottingham, UK, Jul 1999, pp. 511- 527.[2] T. Himmer, T. Nakagawa, M.

120-128.[6] J. Park, M. J. Tari, and H. T. Hahn, "Characterization of the Laminated Object Manufacturing(LOM) process," Rapid Prototyping Journal, Vol. 6, No.1, 2000, pp. 36-49.[7] C. Chen, 1996, "Process Insight About LOM Systems," Solid Freeform Fabrication Symp.,Austin, Texas, Aug 1996, pp. 515-522.[8] S. S. Pak, and G. Nisnevich, "Interlaminate Strength and Processing Efficiency Improvements inLaminated Object Manufacturing," 5th International Conference on Rapid Prototyping, Dayton, Ohio, 1994, pp, 171-180.[9] L. Flach, D. A. Klosterman, and R. P. Chartoff, "A Thermal Model for Laminated ObjectManufacturing Process," Solid Freeform Fabrication Symp., Austin, Texas, 1997, pp. 677-688.[10] L. Flach, M. A. Jacobs, D. A. Klosterman, and R. P. Chartoff, "Simulation of Laminated ObjectManufactring (LOM) with Variation of Process Parameters," Solid Freeform Fabrication Symp., Austin, Texas, 1998, pp. 407-416.[11] F. O. Sonmez and H. T.

Chartoff, "Simulation of Laminated ObjectManufactring (LOM) with Variation of Process Parameters," Solid Freeform Fabrication Symp., Austin, Texas, 1998, pp. 407-416.[11] F. O. Sonmez and H. T. Hahn, "Thermomechanical Analysis of the Laminated Object Manufacturing(LOM) Process," Rapid Prototyping Journal, Vol. 4, No. 1, 1998, pp 26-36.509.

vanced Ceramic Materials and Processes for Three-Dimensional Printing (3DP)Ben Utela,1 Rhonda L. Anderson,*1 and Howard Kuhn21 University of Washington, Dept. of Mechanical Engineering, Box 352600, Seattle, WA 98195-2600, USA 2The Ex One Company, 8001 Pennsylvania Avenue, Irwin, PA 15642, USA. *Author to whom correspondence should be addressed (anderson@me.washington.edu)Abstract: The University of Washington and ExOne, Inc. are collaborating in the development of advanced ceramic materials and processes for three-dimensional printing (3DP). The focus of the research to be presented is work funded by the National Science Foundation to develop a biocompatible alumina-based system for medical and dental applications. Materials design, characterization, and processing considerations will be discussed.290I.SUMMARY OF THE RESEARCH CONDUCTEDThis research focused on the development of a novel class of versatile, high-yield nanoparticle ceramic systems.

In particular, these systems are intended to be compatible with the Three-Dimensional Printing (3DPTM) process for fabrication of dental restorations. In addition to this specific application, the research was broadly targeted to impact other structural ceramic applications through 3DP, and to provide a vehicle for the training of graduate students in this interdisciplinary area of significant commercial interest.I.1 Motivation for the Research, Objectives and Specific TasksThe present research is motivated by a growing market opportunity for ceramic dental restorations via 3DP, on the heels of the current successful commercialization of 3DP application to metal-based dental restorations. Primary research objectives, then, were to determine the feasibility of producing structurally sound ceramic composites by introducing suitable pre-ceramic polymers into the material through the printhead nozzles and/or the print bed in the 3DP process.

To meet these objectives, the research tasks included: (i) identification and characterization of potentially suitable print bed ceramics, pre-ceramic polymer binders, and nanoparticle additions; (ii) evaluation of binder compatibility with 3DP printheads; (iii) evaluation of print bed ceramics containing nanoparticles; (iv) thermal processing of fabricated ceramic parts. In all cases, the evaluations were carried out within the constraints imposed by the 3DP process; specifically, these constraints involve limits on physical properties of the binder (viscosity, density, surface tension) and the print bed ceramic powders (flowability).I.2ResultsMaterials Characterization After screening various candidates, initial characterization focused on Acetate Alumoxane (AA) in both liquid and dry form as a key source of nano particles to be evaluated for this research.

A solution of AA was provided as a 10% volume solution in H2O by Oxane Materials, Houston, TX for its evaluation as a potential 3DP ceramic material. As received, the AA solution was a thick gel with high turbidity. Surface tension was measured at 62.5 dynes/cm ((cid:114) 20%) and viscosity at 17.64 cPs ((cid:114) 2.5%).For compatibility with the R-1 3DPrinter printhead, surface tension of the fluid must be in the range of 25 to 50 dynes/cm and viscosity between 1.5 and 20 cPs. While viscosity of the AA gel is on the upper end of printability, surface tension is out of range. Particle size analysis performed by Spectrex (Figure 1) shows an average suspended particle size of 2.11 (cid:541)m with a standard deviation of 1.28 (cid:541)m. Such particle size is too large to be jetted through the drop-on-demand printhead of the R-1 3DPrinter.

Such particle size is too large to be jetted through the drop-on-demand printhead of the R-1 3DPrinter. Nevertheless, while this source of AA is not suitable for printing, it may provide a source of near nano-sized particles in an additive to printed ceramic materials for enhanced green strength or accelerated sintering.291Figure 1: Results of AA particle size analysis performed by SpectrexSimilarly, screening of various ceramic powders for the print bed in 3DP led to selection of Sumitomo AA-18 alumina. (In this designation, AA identifies the manufacturer’s particular grade of alumina, and is not to be confused with the AA acronym for Acetate Alumoxane.) Due to the requirements of powder spreadability for successful 3DPrinting, it was necessary to find an alumina powder that flows well without agglomerating. To satisfy these parameters the particles should, ideally, be spherical and have relatively large particle sizes on the order of 15 – 30 (cid:541)m.

To satisfy these parameters the particles should, ideally, be spherical and have relatively large particle sizes on the order of 15 – 30 (cid:541)m. The Sumitomo AA-18 powder, Figure 2, is of high quality, has a particle size range of 18 – 26 (cid:541)m, and is made up of faceted, equiangular particles. Particle size analysis is shown in Figure 3. The mean particle size of the powder is 20.93 (cid:541)m (SD 1.298) and the median particle size is 20.95 (cid:541)m. Ninety percent of the powder is between 11.5 (cid:541)m and 31.4 (cid:541)m (5th – 95th percentile). Visual examination of the bulk powder and spreading trials showed that the powder flowed well and created a smooth, even powder layer in the print bed of the R-1 3DPrinter.

Visual examination of the bulk powder and spreading trials showed that the powder flowed well and created a smooth, even powder layer in the print bed of the R-1 3DPrinter. Bed packing density ranges from 50 – 55% depending on the spreading parameters and the layer thickness.292Figure 2: SEM image of Sumitomo AA-18 alumina powder, 500x magnificationFigure 3: Particle size analysis of Sumitomo AA-18 alumina performed using a Micromeritics Saturn DigiSizer 5200AA solution as a liquid binder One use of the AA is as an addition to ceramic print bed powders, injected via the 3DP printhead. The role of the AA in this case would be to increase green strength and increase the ceramic yield in the fired material.

The role of the AA in this case would be to increase green strength and increase the ceramic yield in the fired material. Suitable mixtures of alumina powder and293AA solution were identified for two ranges of particle sizes potentially suitable for 3DP: 1) Sumitomo alumina powder, particle size range 18 – 26 (cid:541)m with a mean particle size of 21 (cid:541)m; 2) Aldrich alumina powder with particle size <10 (cid:541)m.Cast bars (5 x 10 x 25 mm) using AA solution as the binder in combination with the Sumitomo alumina powder were prepared, using a ratio of 7 parts ceramic powder to 1 part AA by weight. The green cast specimens were relatively durable and could be handled without breaking. The castings break cleanly as opposed to crumbling, which is a good indicator of cohesive green strength.These results indicate that green parts with acceptable green strengths are possible using the selected materials.

However, due to the large particle size of the Sumitomo alumina, these samples did not sinter to the desired density, even with the most aggressive sintering run – 6 hours at 1650 ˚C.AA solid as a liquid activated binder Another use of AA is as an additive in solid particle form to the print bed ceramic powders. In the solid state, AA is a fine white powder with particle sizes dependent on the liquid processing parameters. On wetting and subsequent drying the powder is observed to clump and bind together. This observation led to the idea of using dry AA as a liquid activated in-bed binder.Dry AA was provided as a large grain crystalline powder with some particles as large as 5 mm in diameter. This powder was ball milled in grinding media for 24+ hours until it reached the form of a fine powder, shown in Figure 4. Particle size range of this powder was measured to be 1 – 40 (cid:541)m, which is acceptable for spreading by the 3DPrinter.

Particle size range of this powder was measured to be 1 – 40 (cid:541)m, which is acceptable for spreading by the 3DPrinter. While a powder bed of pure AA powder could likely be printed successfully, the ceramic yield of any such printing would be around 60% and significant shrinkage would be observed.Instead, a mixture of AA powder and alumina powder would be desirable as it provides a liquid activated binder in the powder bed, but retains the high yield of a nearly pure ceramic powder bed. Powdered AA in the bed also allows a variety of choices for the activating liquid. This approach has excellent flexibility and the potential to tailor the properties of the raw materials, green part, and ceramic product. Both distilled water and a mixture of water and isopropanol have been found to work as activating liquids. Isopropanol does not activate the alumoxane effectively, but does act to lower the surface tension.

Isopropanol does not activate the alumoxane effectively, but does act to lower the surface tension. Using the AA as an in-bed binder minimizes many of the binder issues such as printhead clogging, waste disposal, cost, and environmental considerations associated with transmission of AA through the printhead.294Figure 4: SEM image of powdered alumoxane, 2000x magnificationTests using powdered alumoxane as the in-bed binder indicated only a very small decrease in the bed’s ceramic yield after firing. TGA and binder elimination studies showed the powdered alumoxane has a 60% ceramic yield by weight (Figure 5).Figure 5: TGA results of pure alumoxane295As described in the previous section, only a small amount of AA gel (one part in 7, by weight) is necessary to create suitable green parts. Experimental studies were performed to evaluate the binding ability of 2.5%, 5%, 7.5%, 10%, and 15% weight mixtures of dry AA with Sumitomo AA-18 alumina.

Experimental studies were performed to evaluate the binding ability of 2.5%, 5%, 7.5%, 10%, and 15% weight mixtures of dry AA with Sumitomo AA-18 alumina. The optimum weight AA mixture with Sumitomo alumina and water binder was determined to be 5%. This mixture resulted in a ceramic yield of ~98.5%, based on TGA analysis (Figure 6). Bars cast with more than 5% alumoxane were slightly better in strength, but resulted in lower ceramic content; bars with less that 5% did not have sufficient green strength. Taking this as the optimum balance between green strength and ceramic yield, a batch of 5% alumoxane/95% Sumitomo ceramic powder was prepared for printing tests.Figure 6: TGA of sample from 5% alumoxane/95% Sumitomo powder mixtureSintering studies In order to determine the sinterability of various biocompatible ceramics, a range of alumina powders and their mixtures with zirconia were cast together with the powdered alumoxane as the binder.

Green and sintered densities, as measured by the Archimedes technique, are shown in Table 2. As expected, finer particles do improve the sintered densities and it is possible to obtain relatively high sintered density of approximately 75%, even with the large alumina particles that are required for spreading in the 3DPrinter.Table 2: Measured densities of Sintering Test MixturesSTM #Density (% theoretical) GreenFired 1570 C (1 hr)Fired 1675 C (1 hr)1 2 3 4 5 6 7 8 9 10 11 12 13 14 1563.71% 63.79% 63.37% 60.42% 61.68% 66.68% 63.46% 66.30% 67.55% 59.36% 61.22% 59.54% 52.39% 46.02% 45.69%67.54% 68.35% * 63.23% 65.44% 71.36% 70.62% 77.64% 73.23% * 65.24% 65.10% 57.69% 53.25% 52.71%66.86% 66.25% 69.08% 63.39% 68.04% 70.90% 66.44% 72.45% 74.61% 61.62% 64.50% 61.45% 62.97% 54.49% 53.16%Wax coating infiltrated porous sample and produced unreliable dataSEM images of STM samples #3, #6, and #9 fired at 1675 ˚C are shown in Figures 7 – 9, respectively.

The large particles are the Sumitomo ceramic powder while the smaller particles are a combination of the additive powder and the AA powder. As seen in the higher magnification images, a limited amount of sintering occurs between the larger alumina particles. It is clear that the finer ceramic particles act as binder with some sintering between the larger particles.297Binder7.6% 7.9% 7.3% 8.3% 7.5% 7.5% 7.6% 7.6% 7.5% 7.7% 7.5% 7.6% 7.5% 7.5% 7.8%Figure 7: SEM images of bulk Sumitomo AA-18 alumina with 22.0% <10 micron Aldrich alumina (STM #3)Figure 8: 1500x and 5000x magnification SEM images of bulk Sumitomo AA-18 alumina with 21.4% InCeram alumina (STM #6)Figure 9: SEM images of bulk Sumitomo AA-18 alumina with 21.6% <5 micron Aldrich zirconia (STM #9)2983Dprinting of Ceramics Two different 3DPrinting methodologies were tested. The more novel system involved the use of powdered AA as an in-bed binder combined with a wetting agent distributed through the printhead.

The more novel system involved the use of powdered AA as an in-bed binder combined with a wetting agent distributed through the printhead. The traditional method used of the ProMetal S-4 binder (developed for stainless steel printing) with a print bed containing a pure ceramic powder. Both methods proved viable.While the results of the sintering tests showed that additive powders are beneficial in increasing the fired density of the ceramic, the small particle additives or ‘fines’ can adversely affect powder spreading. Each of the STM powders that resulted in the highest density samples have more than 20% additive fine powders, which do not spread well. Pits and cracks are formed in the bed as each new layer is spread, and no combinations of layer thickness, roller speed, or carriage traverse speed were found to avoid these issues.

Pits and cracks are formed in the bed as each new layer is spread, and no combinations of layer thickness, roller speed, or carriage traverse speed were found to avoid these issues. A combination similar to STM #5 containing 18.5% InCeram alumina, 78.6% Sumitomo alumina, and 2.8% alumoxane was found spreadable, but there were limitations on the spreading parameters that would result in an acceptable layer spread. Other attempts at spreading Sumitomo alumina powder with 10-20% fines were performed, but all required spreading optimization. Given that a mixture of 95% Sumitomo alumina and 5% AA powder proves spreadable with almost any parameters, this mixture was used for the bulk of the printing tests. Use of this mixture also reduces the powder bed from a three-part to a two-part mixture and helps to reduce the number of variables that could affect the 3DP parts.

Use of this mixture also reduces the powder bed from a three-part to a two-part mixture and helps to reduce the number of variables that could affect the 3DP parts. Table 3 provides initial 3DP parameters that were used to help guide materials and binder development and selection.Table 3: 3DP parameter identifications for successful printing and spreadingBinder FormulationPrint head specs Extreme range ProMetal targetViscosity (cps) 5 - 15 1.5 - 20 7.5Surface Tension (dynes/cm2) 28 - 35 25 - 50 30 - 32Density (g/cm3) none none 0.9 - 1.35Powder FormulationTypical powder S4D, high dens Suggested idealParticle Range (cid:541)m 10 - 80 10 - 80, sm. 5 - 10 15 - 20Shape Spherical Spherical SphericalDensityBimodal No Yes No, small dist.Since the selected powder mixture could be activated with water it was decided that initial testing be done with distilled water.

A trial-and-error approach allowed the creation of a repeatable droplet mass by weighing the drops created by different piezoelectric waveforms. After optimization of the saturation levels it was possible to successfully print parts with reasonable surface feature definition. While the distilled water proved to be a workable system, it was decided to develop a new binder with surface tension closer to the ranged targeted for the ProMetal printhead of 28 – 35 dynes/cm2 in hopes of easing the optimization of the printing process. A mixture of 80% H2O/20% Isopropyl alcohol, having a surface tension of roughly 35 dynes/cm2, met this objective. This new binder also has a viscosity of 1.81 cPs as opposed to 0.99 cPs for299Lubricating powder 0.25% SiO20.25% SiO2distilled H2O.

This new binder also has a viscosity of 1.81 cPs as opposed to 0.99 cPs for299Lubricating powder 0.25% SiO20.25% SiO2distilled H2O. While this viscosity is still lower than the printhead specifications, the alcohol/ H2O mixture appeared to be a more reliable binder solution.Demonstration parts using AA in the powder bed are shown in Figure 10. Considerable definition is observed in the letters, but the block corners are typically smoothed during manual handling.Figure 10: 3DP blocks printed using H20/Isopropal binder with alumoxane/alumina powder bedThe results of printing the ProMetal binder into the Sumitomo alumina are extremely impressive. Figure 11 shows a test array of 25 x 25 x 6 mm block with 3 mm walls and posts. These two blocks were printed at right angles to each other so one was built with the walls in line with the path of the printhead, and the other with the walls perpendicular to it. The orientation causes significant differences in the parts.

The orientation causes significant differences in the parts. The parallel array prints the slots better, but the perpendicular array prints a very thin post successfully.Figure 11: Test arrays printed with ProMetal binder and Sumitomo AA-18 powder, parts parallel and perpendicular to the print direction, respectively300Additional parts were printed using the ProMetal binder and alumina powder to demonstrate the feasibility of producing dental copings and crowns, one direct intended application of this research. Copings are shown in Figure 12 and full crowns in Figure 13.Figure 12: Green copings printed with ProMetal binder and Sumitomo AA-18 aluminaFigure 13: Fired molars (1575 C) printed using ProMetal binder and Sumitomo AA-18 powder. Approximate dimensions are 8 x 8 x 8 mm.301Given the low sintered densities of the various ceramic powder mixtures, infiltration of the printed parts with glass.

Approximate dimensions are 8 x 8 x 8 mm.301Given the low sintered densities of the various ceramic powder mixtures, infiltration of the printed parts with glass. Preliminary experiments were performed using In-Ceram Z21 Zirconia Glass Powder and the Sumitomo AA-18 alumina molars shown in Figure 13. Using the procedure recommended by In-Ceram, a slurry of water and the glass powder infiltrant was mixed and brushed onto the sample. As the water dried the infiltrant adhered to the sample. The coated part were then fired to 1150 ˚C, which led to successful infiltration of the exterior of the part. The maximum penetration of the glass was about 1.5 mm on a molar with an eight-millimeter diameter. Figure 14 shows a SEM of a cross section of the infiltrated molar. The top of the image shows the infiltrated region.

Figure 14 shows a SEM of a cross section of the infiltrated molar. The top of the image shows the infiltrated region. Although the infiltrant only entered to a depth of 1.5 mm, this is sufficient for full penetration in the thin-walled copings shown in Figure 13.Figure 7: SEM image of infiltrated tooth cross section, upper section infiltrated with glass, lower section uninfiltrated, 500x magnificationWhile only in the proof of concept stages, the test infiltration of ceramic printed parts shows great promise for the production of fully dense ceramic-matrix materials of customizable properties.302IV.ConclusionsThis research was successful in developing a biocompatible ceramic composite with commercialization potential for rapid manufacturing by 3DP.

The principles learned about processing parameters, thermal conversion, and final properties of the simple printed shapes are essential for the robust and reproducible processing of complex ceramic parts by solid freeform fabrication methods, and will be applied to geometrically intricate parts in a planned future effort. Overall, the results of our research into ceramic materials compatible with the 3DP process will accelerate the evolution of 3DP methods from prototyping to manufacturing as well as enhance our opportunity for timely entry into the commercial marketplace.In addition to the significant impact on broader inkjet and 3D printing research and commercial opportunity, this project involved the academic and career preparation of students at many levels.

Graduate students and post-doctoral researchers worked closely with the academic and industrial team in aspects of materials synthesis, processing and characterization, and had the opportunity to conduct a portion of their research onsite with ExOne scientists and engineers. Furthermore, plans are underway to expand research programs in 3DP with a broader range of biocompatible materials as well as to incorporate these areas into three multidisciplinary courses currently offered at the University of Washington, across multiple departments. Research in this way will continue to offer a broad, real-world educational opportunity in cutting-edge technology development for students as well as academic and industrial researchers.The authors gratefully acknowledge support for this research from the National Science Foundation, under Award Number NSF DMI-0512682.303.

vanced Heat Sinks Enabled by Three-Dimensional PrintingAlan Lyons1,2*, Shankar Krishnan1,3, John Mullins1, Marc Hodes1,4, and Domhnaill Hernon11Bell Laboratories – Alcatel-Lucent Dublin, Ireland2Current Address: CUNY – College of Staten Island New York, NY, U.S.A. Lyons@mail.csi.cuny.edu3Current Address: Battelle/Pacific Northwest National Laboratory, MicroProducts Breakthrough Institute, Corvallis, OR4 Current Address: Tufts University, Medford, MA, U.S.A.Abstract:With the rapid rise in power dissipated by integrated circuits, improved heat sinks designsare needed to decrease the thermal resistance between them and forced air streams. Manufacturing methods such as extrusion, machining and die-casting have been used to fabricate conventional longitudinal fin designs. Although these technologies add relatively little cost, they preclude the fabrication of more complex heat sink designs.

Although these technologies add relatively little cost, they preclude the fabrication of more complex heat sink designs. We discuss novel heat sink designs which increase surface area and/or modulate air flow streams. Fabrication of these unconventional designs is enabled by using 3D printing technologies with the subsequent conversion of the printed parts into monolithic copper structures by investment casting.Introduction:The use of longitudinally finned heat sinks in electronics cooling is ubiquitous. With everincreasing thermal loads and densities such conventional heat sink designs, with 2-dimensional flat fins, have reached the limit of their usefulness for some applications and will not be capable of providing adequate cooling for future high density and functionality products.

For example, some circuit packs dissipate more than 300 Watts of power and thermal designers are struggling to accommodate such high heat dissipation rates using conventional, longitudinally finned heat sinks while maintaining junction temperatures below those required to assure the long term reliability requirements of telecommunications products. Often, the only viable means of insuring adequate heat dissipation is by attaching the heat sink bases to a vapor chamber to enhance heat spreading over larger areas. This solution adds considerable complexity and cost to the circuit packs.One approach to increasing the heat dissipating capacity of 2D heat sinks is to increasethe velocity of the forced air flow through them. Higher air flows will increase the heat transfer coefficient by decreasing the static boundary layer thickness between fin and air and, moreover, decrease the caloric temperature rise of the air.

Higher air flows will increase the heat transfer coefficient by decreasing the static boundary layer thickness between fin and air and, moreover, decrease the caloric temperature rise of the air. Higher levels of acoustic noise and pressure drop, however, accompany increased air flow rates. For many applications these acoustic limits (e.g., 85 dBa for data center products) have already been reached and consumer electronics are strivingCorresponding authorto reduce noise levels. In addition, faster air flow rates require greater electrical power to achieve, thus increasing the overall power consumption and carbon footprint of the product. In short, increasing the air flow rate can increases the capacity of 2D heat sinks, but, beyond a certain limit, this approach is unfeasible due to both acoustic noise and power consumption limits.A second approach that can be used to increase heat sink performance is the use of highsurface area fins.

For example, non-planar fin arrays have been commercially available for many years (e.g., elliptical pin-fin arrays, bent folded fins, etc.). However both the magnitude of the deviation from a flat fin, as well as the performance enhancement achieved by such designs, is limited. One reason for this small impact is that the overall surface area of such designs is not significantly different than for flat fin heat sinks. Metal foams exhibit a significantly increased surface area as compared to these designs. Sandwiching stochastic metal foams between flat fin heat sinks has been reported [1] where an adhesive was used to bond the foam to the fin.

Sandwiching stochastic metal foams between flat fin heat sinks has been reported [1] where an adhesive was used to bond the foam to the fin. These heat sinks have not found practical applications, however, as their performance is limited by two effects: (1) a low thermal conductivity interface between fin and foam resulting from the use of an adhesive to assemble the components and (2) a significant pressure drop across the heat sink resulting from the tortuous air flow path created by the random size and alignment of pores in the foam.Other approaches to increase heat transfer have been demonstrated by incorporating structures that introduce beneficial flow characteristics including flow unsteadiness or flow instabilities (such as Tollmien-Schlichting waves or Kelvin-Helmholtz instabilities). Patera and Mikic [2] introduced the concept of resonant heat transfer enhancement based on excitation of shear-layer instabilities in internal separated flows.

Patera and Mikic [2] introduced the concept of resonant heat transfer enhancement based on excitation of shear-layer instabilities in internal separated flows. They reported three to four fold increase in heat transfer augmentation under given viscous dissipation penalties over conventional channel geometries [3]. It was shown by the Reynold’s analogy that the problem of designing optimal heat-transfer enhancement systems is best considered as a problem in hydrodynamic stability theory: a more unstable flow will generate larger Reynolds fluxes at lower Reynolds numbers, and thus achieve commensurate heat transfer at a fractional dissipative penalty [2].

In order to exploit the natural instabilities Patera and Mikic proposed the following system requirements [2]: creation of a system with separated flow; determination of the system’s resonant frequency; and excitation of that frequency with appropriate modulation.Conventional heat sink manufacturing processes, however, are limited in their ability tocreate structures that can increase surface area, induce flow unsteadiness, or modify air flow patterns in some other, useful way. Extrusion is the most widely used technique to manufacture heat sinks as it is especially well suited to create a monolithic component composed of parallel fins on a heat-spreading base plate. This process has many advantages including low cost, simple tooling, limited post-extrusion processing and fins monolithic with the base of the heat sink. With a simple 2D die, thin fins of aluminum can be formed; typically less than 1.0mm thick.

With a simple 2D die, thin fins of aluminum can be formed; typically less than 1.0mm thick. By partially cutting the extrusion perpendicular to the fins, a square pin-grid array of fins can be formed which may increase surface area slightly. Inherent to extrusion, however, is the formation of parallel flat fins as only a 2D die is possible.To achieve more complex fin designs other manufacturing processes must be employed. Stamping and assembly processes are used to fabricate closed channel heat sinks for today’s high power microprocessor chips. This approach enables the use of thinner fins (<0.5mm) as well as the formation of closed (i.e., ducted) channels which increase surface area for a given heat sink volume. Fins are formed from stamped metal into a “C” shape, mechanically interlocked into parallel channels and soldered onto a heat-spreading base.

Fins are formed from stamped metal into a “C” shape, mechanically interlocked into parallel channels and soldered onto a heat-spreading base. However, these assembled heat sinks remain essentially parallel finned structures.Lost wax casting is a process which enables metal heat sinks to be manufactured with significantly more detail and a greater variety of shapes. The original part (fabricated by any technique including machining metal) is used to create a rubber mold from which multiple wax patterns are formed. The wax patterns are connected by sprues and surrounded by a plaster-like investment. By heating the investment, the wax melts away leaving a replica in the investment which is subsequently filled with molten metal. After removing the investment, faithful duplicates of the original part are obtained. The only requirement to replicate a heat sink by investment casting is that the shape can be removed from the mold.

The only requirement to replicate a heat sink by investment casting is that the shape can be removed from the mold. For example, pin fin arrays where each pin has an elliptical cross-section can be fabricated by such lost wax casting techniques. Many other shapes (e.g., re-entrant features), however, cannot be molded as they incorporate features which cannot be separated from the mold and thus currently find no suitable path for manufacture.In this paper, we describe a sacrificial pattern casting technique which enables the manufacture of essentially any 3D shape as a monolithic heat sink. A commercially available Multi-Jet Modeling (MJM) system is used to fabricate the original heat sink models. These models are then used as sacrificial patterns in an investment casting process. By combining these two techniques, any arbitrarily shaped fin, duct or channel can be created.

These models are then used as sacrificial patterns in an investment casting process. By combining these two techniques, any arbitrarily shaped fin, duct or channel can be created. Complex, 3D designs can be cast as a monolithic structure, eliminating the thermal contact resistance created when solid surfaces are attached by conventional means (e.g., adhesively bonded or soldered). With fabrication constraints removed, the challenge becomes the design of a heat sink which exploits this new fabrication freedom to improve thermal performance. We present three novel heat sink design approaches, enabled by the sacrificial pattern casting process, which increase the heat that can be transferred to a forced convection air stream.Experimental:Computer aided design (CAD) files of the heat sink designs were created using ProEngineer or 3dsMax software and exported as .stl files.

These files were sent to an InVision® HR 3-D Modeler, manufactured by 3D Systems, which is a commercially available MJM system. In the InVision HR, heat is used to reduce the viscosity and facilitate jet dispensing of the wax based support material (VisiJet® S100) and the pre-polymer model material (VisiJet® HR 200). The latter is cured upon exposure to UV light after completion of one or more build layers. After completion of the build, support material was removed by heating the part in a forced air convection oven at 70oC. Casting was performed by Best-Cast (River Edge, NJ) where sprues were attached to the pattern before being encased with investment. The pattern was burnt-out from the investment at high temperature leaving a cavity mold within the hardened investment. The investment was inserting into a vacuum casting machine and evacuated to remove trapped air. Then, molten metal was forced into the mold using pressurizedinert gas.

The investment was inserting into a vacuum casting machine and evacuated to remove trapped air. Then, molten metal was forced into the mold using pressurizedinert gas. A copper alloy was used to reduce the casting temperature and improve flow of the molten metal. After casting, the investment material was removed with a water jet and sprues were cut and sanded.To facilitate testing and minimize thermal interfaces, the heat sinks were incorporatedinto a test platform to hold them in proper position within a wind tunnel and provide a mounting location for a thin-film heater. The test platform and heat sink were cast together as a unified, monolithic structure as shown in Figure 1. The heat sink volume (see Figure 1) measures 32 mm x 15 mm x 32 mm (L x H x W). The heat sink is integrated with a flanged base. The base (32 mm x 15mm x 32mm) is hollow to accommodate a thin-film heater (Minco).

The heat sink is integrated with a flanged base. The base (32 mm x 15mm x 32mm) is hollow to accommodate a thin-film heater (Minco). The flange facilitates mounting within the wind tunnel.Figure 1 Fin-foam heat sink (7 fins) cast with base and mounting flange in a copper alloy as seen in perspective and instrumented for wind tunnel measurements (a) and from the back side (b).The governing flow and heat transfer equations for incompressible, steady flow of a Newtonian fluid are solved using the commercial software FLUENTTM. The computational geometry was created using the software GAMBITTM. The geometry was meshed using hybrid (tetrahedral and hexagonal) elements in GAMBIT by specifying the minimum edge length. The mesh so created was exported to FLUENT for fluid and heat flow simulations. A second-order upwind scheme was used for the flow and heat transfer calculations.

The mesh so created was exported to FLUENT for fluid and heat flow simulations. A second-order upwind scheme was used for the flow and heat transfer calculations. A co-located pressure- velocity formulation in conjunction with the SIMPLE algorithm was used for obtaining the velocity fields, and the linearized systems of equations are solved using an algebraic multigrid algorithm. For the calculations reported here, a total of approximately 1 million finite volumes was found to be sufficient and were used for numerical calculations.The wind tunnel used to characterize the heat sinks is an ATS CWT-100 and consists ofhoneycomb, screen, contraction and screen sections, respectively, upstream of the test section inlet to reduce the background turbulence intensity of the flow and to produce a uniform velocity profile in the test section. The test section is made from Plexiglas and the prototype heat sinks were placed in a fully ducted arrangement within the wind tunnel.

The test section is made from Plexiglas and the prototype heat sinks were placed in a fully ducted arrangement within the wind tunnel. The wind tunnel is powered by two 12W fans that are placed downstream of the diffuser section. Wall mounted static pressure taps, located upstream and downstream of the heat sink, were used for pressure drop measurements. Type-T thermocouples were used to measure the inlet air and maximum heat sink temperatures. A Minco heater with pressure sensitive adhesive was affixed to the base of the heatsink to simulate microprocessor power input. To mitigate heat losses to the environment a Aspen aerogel foam insert was placed directly on the back-side of the heater. Inlet velocity to the heat sink was measured (approximately 30 mm) upstream of the heat sink in the test section of the wind tunnel using a Pitot-static probe (United Sensors).

Inlet velocity to the heat sink was measured (approximately 30 mm) upstream of the heat sink in the test section of the wind tunnel using a Pitot-static probe (United Sensors). The heat sinks were characterized in the wind tunnel by measuring the thermal resistance (i.e., ratio of difference between maximum temperature of the heat sink and inlet air temperature to heater power) of the heat sink against pressure drop, pumping power and velocity.Results and Discussion:Heat Sink Design Approaches. The objective of any air-cooled heat sink design is to optimize the distribution of highly thermally conductive material in the allowable volume to minimize thermal and flow resistances. In general, any increase in surface area increases the heat dissipated from the surface. Concomitantly, however, fluid frictional loss increases thereby increasing the pressure drop across and caloric resistance of the heat sink.

Concomitantly, however, fluid frictional loss increases thereby increasing the pressure drop across and caloric resistance of the heat sink. The challenge is to optimize the heat sink design to increase the heat transfer with minimal fluid frictional losses. From a conceptual framework, enhancement of heat transfer can be achieved under minimal fluid frictional losses by exploiting various heat and fluid flow phenomena such as boundary- layer restarting, tripping of boundary layers and self-sustained flow unsteadiness [4].Our goal was to demonstrate the usefulness of the sacrificial pattern casting approach tofabricate heat sinks with enhanced performance by incorporating such air flow manipulation features into the structure.

A three-pronged approach was undertaken to enhance heat transfer from a wall by streaming the flow through a two-dimensional array of polygonal ducts (Hexagonal Perforated Duct Arrays), by introducing flow-obstacle-induced local mixing (Ordered Metal Foams), and by exploiting hydrodynamic instabilities to sustain flow unsteadiness (Schwartz Structures).Hexagonal Perforated Duct Arrays. Ducted heat sinks are one of the simplest ways toincrease surface area compared to flat, 2D fin heat sink designs. By enclosing the channel, more metal surface is present over which heat may be transferred to the flowing air stream. However, the challenge becomes how to design the duct array to increase surface area without introducing a significant pressure drop into the system. Honeycomb structures have been discussed for heat exchanger applications [5] where the honeycomb structure was brazed or attached via thermal adhesive to the upper and lower heat transfer surfaces.

Honeycomb structures have been discussed for heat exchanger applications [5] where the honeycomb structure was brazed or attached via thermal adhesive to the upper and lower heat transfer surfaces. As stated previously, this will cause extra thermal barriers that will reduce the effectiveness of the design. Also, these honeycomb structures were limited to straight channels and did not contain any perforations that manipulate the flow to increase the heat transfer.To optimize the cooling capacity of a two-dimensional duct arrays, the performance of amodel system was analyzed for various cell shapes and arrangements. The model reported by Gu et al [6] was modified here for heat sinks applications. Details of the model can be obtained from [6]. The heat sink volume considered was 32 x 15 x 32 mm (L x H x W). A constant pressure drop value of 1 Pa was fixed across the heat sinks.

Details of the model can be obtained from [6]. The heat sink volume considered was 32 x 15 x 32 mm (L x H x W). A constant pressure drop value of 1 Pa was fixed across the heat sinks. Fin equations were written in generalized local coordinates for heat transfer from the wall to the air flow through an array of polygonal ducts. Fully developed flow conditions were assumed and local Nusselt number and friction factors were specified for the duct shapes being analyzed. After a series of algebraicmanipulations, an analytical expression for thermal resistance as function of geometry and flow parameters was obtained[7]. Figure 2 plots the thermal resistance as a function of cell pitch, l, for a fixed fin/wall thickness of 0.5 mm.

Figure 2 plots the thermal resistance as a function of cell pitch, l, for a fixed fin/wall thickness of 0.5 mm. As seen clearly from the figure, optimized hexagonal ducts significantly out-performed the other shapes for a fixed pressure drop value across the heat sinks of 1 Pa.Figure 2: Predicted thermal resistance as a function of cell pitch for triangular, square and hexagonal ducts.Having assessed the effect of the shape of the duct, numerical simulations were performed on hexagonal ducts to precisely predict their optimal size [7]. A periodic unit-cell was used and the cell-pitch and cell-thickness was varied systematically based on the initial guesses derived from the analytical solutions.

A periodic unit-cell was used and the cell-pitch and cell-thickness was varied systematically based on the initial guesses derived from the analytical solutions. A 10% improvement in thermal resistance for hexagonal ducts was observed over a conventional parallel plate heat sink.It can be shown by a simple analytical treatment for fully-developed thermal and fluidflow conditions that there exists a critical flow length above which having multiple ducts instead of a straight channel would not be optimal. In order to further improve the performance of cellular structures, periodic slots or perforations were introduced to restart the boundary layers and to mitigate the problem of non-optimal number of ducts. The boundary layer is a region adjacent to the duct wall of relatively static air and containing strong gradients in the flow thereby acting as a thermal insulator from the cooling effect of the free stream air flow.

By disrupting the boundary layer growth and subsequently restarting the boundary layer development, heat transfer will be increased owing to the development of thinner boundary layers. Further, the introduction of slots promotes local mixing of hot air near the heat source with cold air near the top of the heat sinks.One example of an enhanced 3D heat sink that incorporates hexagonal ducts is shown inFigure 3. This is a type of cellular structure in which fluid flows through ducts which are periodically perforated by slits. Although it would be possible to manufacture continuous hexagonal ducts by extrusion, such extrusion die would be complex and expensive to fabricate, necessitating a large volume application to recover the development and tooling costs. Perforation of the hexagonal ducts with slits would not be possible by extrusion or conventional casting techniques alone. Three-dimensional printing allows for the introduction of slots anywhere along some or all of the ducts.

Three-dimensional printing allows for the introduction of slots anywhere along some or all of the ducts. The hexagonal perforated duct heat sink fabricated by the sacrificial pattern casting process has several advantages over both hexagonal duct and parallel fin heat sinks. First, as the heat sink is monolithic, no thermal barriers exist between hexagonal ducts or between the base and the ducts. The heat transfer surface area has increased substantially over a parallel fin heat sink with the same volume. In addition, the honeycomb channels can be continuous or they can incorporate openings of any design in either the horizontal and/or the vertical directions.Figure 3: CAD model of a honeycomb heat sink with slits (a), polymer pattern created with the InVision MJM printer (b) and copper alloy heat sink created by the sacrificial pattern casting process.Ordered Metal Foams.

Metal foam structures can be generated with any arbitrary geometry, either random or ordered, including body-centered cubic (BCC), face-centered cubic (FCC) or A15 lattice arrangements. The foam can be located between fins, as shown in Figures 1a and 4, or the heat sink could be made of just the foam structure and have no fins. One of the key advantages is that a monolithic foam structure can be manufactured, eliminating the need for low thermal conductivity interfaces. Another advantage of this design is the increase in the surface area that is available for heat transfer compared to a standard heat sink design. For example, the surface area available for heat transfer on a fin/foam structure with 4 fins is more than 15% greater than the surface area of a parallel fin heat sink with identical length, height and width dimensions.Figure 4: CAD model of a monolithic four-fin foam heat sink seen from the front (a) and in perspective (b).

The metal foam located between parallel fins was fabricated with an ordered BCC structure.A third key feature of the fin-foam structures is the beneficial flow characteristics thatare induced downstream of the foam ligaments within the fin passages. Flow through metal foam gives rises to local disturbances in velocity due to eddies that are shed in the wake of the flow past the solid fibers [8]. The presence of pore-level temperature and velocity gradients results in the enhancement of diffusion (or dispersion) sometimes referred to as anomalous diffusion. This enhanced diffusion is often characterized by the total thermal diffusivity. In addition to these thermal dispersion effects, flow instabilities, unsteady, laminar, transitional and turbulent flows can also be set up which increase the heat transfer.Figure 5 shows a typical computational simulation result for modeling fin-foam heat sinks and exemplifies the complexity of the air flow patterns in such structures.

This figure illustrates the predicted temperature distribution in a 4-finned foam heat sink geometry. In order to reduce prohibitive computational cost, only a unit-cell is modeled by exploiting the periodicity in the geometry.Figure 5: Predicted temperature distribution in a 4Finned-Foam Heat Sinks. Inset shows a BCC type periodic unit cell.Figure 6 compares numerically predicted fin-foam heat sink results with the experimentally measured thermal performance of cast heat sinks. Considering the complexity of the geometry of the heat sinks, predicted results are in excellent agreement with experimental values. One exception to this agreement occurs for the metal foam heat sink without fins. In this case, the numerical model imposes a geometrical constraint which restricts transverse dispersion. For fin- foam heat sinks, this constraint is valid as the solid fins physically restrict air flow to the direction along the axial flow axis and prevent transversal flow disturbances.

For fin- foam heat sinks, this constraint is valid as the solid fins physically restrict air flow to the direction along the axial flow axis and prevent transversal flow disturbances. The physical validity of the numerical model results in excellent correlations to experimental measurements. Without fins, air can flow both along and transverse to the axial flow axis, making the numerical model invalid for this case.Figure 6. Comparison of numerically predicted fin-foam heat sink performance with experimentally measured thermal resistance as a function of pressure drop.Schwartz Structures. The concept of excitation of shear-layer (laminar) instabilities for heat transfer augmentation was investigated by fabricating heat sinks based on Schwartz P type minimal surfaces. A minimal surface is a surface with a mean curvature of zero and is locally area minimizing.

A minimal surface is a surface with a mean curvature of zero and is locally area minimizing. In order to exploit the flow instabilities, the geometry of the heat sink needs to have a separated flow regime; then, at the system’s resonant frequency heat transfer augmentation will be achieved. The Schwartz structure, owing to its undulating path, naturally establishes a separated flow and, at appropriate flow velocities, heat transfer augmentation could be achieved. Schwartz surface unit cells were formed into various arrays, as shown in Figure 7.Figure 7: Arrays of Schwartz P-surface unit cells: CAD model of a 3 x 3 x 3 array (a), same model as (a) viewed in perspective (b), copper casting of a 7 x 3 x 7 array (c), copper casting of a 4 x 1.5 x 4 array (d).Thermal and Fluidic Performance of Heat Sink Designs.

Figure 8 shows plots of the thermal resistance (Figure 8a) and pressure drop (Figure 8b) of the three heat sink concepts discussed in this work as a function of air flow velocity. Results from a conventional parallel fin heat sink, fabricated by the same sacrificial casting process, are included in these plots for comparison. It is readily evident that all three novel heat sinks show superior thermal performance compared to a parallel plate heat sink. However these novel heat sink designs incur higher pressure drops, as compared to the parallel plate designs.Figure 8 (a): Experimentally measured thermal resistance of different heat sink concepts compared against a conventional parallel plate heat sink, (b) Measured pressure drop as function of velocity.Conclusion:The new heat sink designs presented in this paper can offer advantages over conventional,longitudinally finned heat sinks, depending upon the application requirements.

For applications where pressure drop is a concern, honeycomb or parallel plate heat sinks will be suitable. For applications where pumping power (which is a more realistic metric because it is a measure of energy expended) is prescribed, any of the three new concepts can be adopted for increasedthermal performance. For applications where pressure drop is not a concern, a finned-foam or Schwartz type structure would be most suitable. Casting these designs from sacrificial polymeric patterns (prepared by three-dimensional printing) is a feasible approach for manufacturing small quantities for specialty products or where conventional heat sink designs are inadequate.

The design space accessible to heat sink designers has been significantly increased by this fabrication approach and we anticipate that further thermal performance enhancements will be forthcoming.Acknowledgements: This work was supported by IDA Ireland.1 Bhattacharya and Mahajan (2002) “Finned Metal Foam Heats Sinks for Electronics Cooling in Forced Convection” ASME J. Elect. Pack. 2. A. T. Patera and B. B. Mikic, “Exploiting hydrodynamic instabilities. Resonant heat transfer enhancement,” International Journal of Heat and Mass Transfer, Vol. 29, pp. 1127 – 1138, 1986. 3. G. Karniadakis, B. B. Mikic and A. T. Patera, “Minimum-dissipation transport enhancement by flow destabilization: Reynolds’ analogy revisited,” Journal of Fluid Mechanics, Vol. 192, pp. 365 – 391, 1988. 4 A. M. Jacobi and R. K. Shah, “Air-side flow and heat transfer in compact heat exchangers: A discussion of enhancement mechanisms,” Heat Transfer Engineering, Vol. 19, pp. 29 – 41, 1998. 5. T. Wen, J. Tian, T.

th Solid Freeform Fabrication Symposium 2004, Austin, TXAdvanced Laminated Object Manufacturing (LOM) of SiSiC CeramicsL. Weisensel, N. Travitzky and P. GreilThe University of Erlangen-Nuremberg, Department of Materials Science,Glass and Ceramics, GermanyAbstractCarbon sheets were used as a starting material for fabrication of SiSiC composites byadvanced LOM. This approach consists of three steps: First a preform was fabricated formphenolic resin coated carbon paper with a LOM-device. Second the preform was turned into acarbon preform by pyrolysis in N2-atmosphere. Third pressureless reactive melt infiltration ofsilicon into the as fabricated carbon preform, which finally yielded a dense SiSiC composite.SEM analysis revealed a microstructure consisting of uniformly dispersed b -SiC grains in amatrix of silicon.

The LOM fabricated material exhibited an average four point bendingstrength and Youngs modulus of 115 MPa and 165 GPa, respectively.IntroductionSiC-based ceramics exhibit superior mechanical properties, high temperature resistance,chemical stability and oxidation resistance from ambient to peek operation temperatures re-quired for advanced applications ranging from automotive to aerospace uses [1-5]. The manu-facturing of SiC-based ceramic composites by reactive infiltration of C/SiC-powder or carbon257fiber reinforced carbon preforms with liquid Si at temperatures between 1450-1600°C in vac-uum (LSI-process) represents an established processing route. Nearly dense interpenetratingor fiber reinforced ceramic composite materials with a variable amount of excessive Si areobtained. Well known materials are REFEL SiC [6] and SILCOMP SiC/Si [7] as well as car-bon fiber reinforced SiC-CMCs [8].

Well known materials are REFEL SiC [6] and SILCOMP SiC/Si [7] as well as car-bon fiber reinforced SiC-CMCs [8]. The LSI-process is also capable for manufacturing nearnet-shape ceramic composites for structural applications [9, 10]. SiC can be manufactured byrapid prototyping technologies, such as selective laser sintering [13], three dimensional print-ing [14] or laminated object manufacturing (LOM) [11,12].LOM produces three-dimensional objects with sheets of paper, polymer, metal or a combi-nation thereof [15]. It can be considered as a hybrid between “subtractive” and “additive”processes: the parts are indeed built up in the layer by layer approach (“additive” processes),but each layer is individually cut by a special tool (“subtractive” processes) in the shape of thecross section of the part.

Each layer is bonded to the previous layer with an adhesive coatingon the bottom side of the sheets (e.g., paper), that is activated by a heat or pressure during theLOM process [15]. A convenient way to fabricate advanced structural ceramics by LOM in-cludes use of ceramic based tapes processed by tape casting. The ceramics obtained includeAl2O3 [16], monolithic ZrO2 and ZrO2/Al2O3 [17], SiC and SiC/SiC [11,12] and Si3N4 [18].The intention of the present work was to demonstrate the fabrication of dense SiSiC laminarcomposites by Laminated Object Manufacturing.Experimental procedureProcessing included preparation of carbon sheets coated with a suitable adhesive,LOM processing and final post-processing with liquid Si infiltration into the porous carboncomponent.

Carbon sheets made from pitch fibers as shown in Fig 1 (Sigratex PE204, Thick-ness ~ 200 µm, SGL Carbon Group, Meitingen, Germany) were laminated to a heat activatedadhesive film containing phenolic resin and butyl nitrile rubber,(Tesa HAF 8410 Tesa AG,258Thickness ~ 60 µm, Hamburg, Germany). After lamination at120°C a carbon sheet coatedwith a phenoly butyl nitrile rubber heat activated adhesive film on one side and a thickness of250 µm was LOM processed in a LOM device (Kira PLT A4, Kira Corporation, Japan). Fig. 2shows the processing scheme applied. Lamination was carried out at 180 °C for 60 s and wasadopted to the curing cycle of the adhesive layer.

Fig. 2shows the processing scheme applied. Lamination was carried out at 180 °C for 60 s and wasadopted to the curing cycle of the adhesive layer. Sample bars which consisted of 20 individ-ual layers each with an overall sample height of 5 mm and a density of 0.42 –0.01 g/cm³were prepared.50 µmFigure 1 Microstructure of the carbon sheets,The porous layered samples were pyrolysed in an electrically heated furnace (GeroGLO 40, Gero, Germany) in nitrogen atmosphere. The heating cycle was optimized usingdifferential scanning calorimetry (DSC) and thermo gravimetric analysis (TGA). A heatingrate of 1 °C min-1 up to 350°C with a holding time of 1 h was applied to avoid cracking andcurling of the sheet during pyrolysis. Further heating up to 800°C was continued with a rate of2 °C min-1. After 1h exposure at 800°C the samples were cooled to room temperature with acooling rate of 5°C/min.

Further heating up to 800°C was continued with a rate of2 °C min-1. After 1h exposure at 800°C the samples were cooled to room temperature with acooling rate of 5°C/min. The porous carbonised samples were then infiltrated with liquid sili-con in vacuum at 1500°C for 1h.259Phenolic Adhesive TapeCarbon SheetsLaminatingLaminated Object Manufacturing(Pressing Temperature 180°C, Time 60s)Pyrolysis (800°C, N 2-Atmosphere)Liquid Silicon Infiltration (1500°C, 1h, Vacuum)Figure 2: Processing scheme.The weight loss and shrinkage upon thermal treatment and post-infiltration were determined.Densities were measured geometrically and by Archimedes method using water as the immer-sion medium. The strut density of the carbonised samples was measured using He-pycnometry analysis. X-Ray diffraction (XRD) analysis (Kristalloflex D 500, Siemens,Karlsruhe, Germany) was conducted using monochromatic Cu-Karadiation at a scan rate of0.75(cid:176) min-1.

X-Ray diffraction (XRD) analysis (Kristalloflex D 500, Siemens,Karlsruhe, Germany) was conducted using monochromatic Cu-Karadiation at a scan rate of0.75(cid:176) min-1. Microstructure and micro chemical composition of the as-pyrolysed and the SiSiCsamples were analyzed by scanning electron microscopy (SEM, Quanta 200, FEI, Germany)equipped with an energy-dispersive spectrometer (EDS, Inca x-sight, Oxford Instruments,UK). The samples for SEM/EDS analyses were ground and polished to a 3 m m diamond fin-ish. At least five different spot analyses were conducted to obtain the average composition ofa given phase.The dependence of fracture strength on the fiber orientation of the fabricated SiSiC bars(3· 4· 45 mm³) was evaluated in four-point bending with the loading direction normal to the260layer extension using spans of 40 and 20 mm according to ASTM standard C1211-92. Thetensile surfaces of the samples were polished to a 3 m m diamond finish prior to loading.

Thetensile surfaces of the samples were polished to a 3 m m diamond finish prior to loading. Thetests were performed using an universal testing machine (Instron 4204, Instron corp. Canton,MA USA) with a crosshead speed of 0.5 mm/min. The Youngs modulus was determined us-ing the impulse excitation technique (Buzz-o-sonic, BuzzMac Software, Glendale, MI, USA)according to ASTM E1876 and C1259.Results and discussionFig. 2 shows the cross section of the LOMed multi layer prior to post-infiltration with Si. Theindividual layers of the carbon sheets and the phenolic butyl nitrile rubber adhesive can bedistinguished in Fig 3a. Fig. 3b shows the microstructure of the pyrolysed body.

Ac-cording to the results of the DSC and TGA measurements the butyl nitrile rubber starts todecompose at 275°C and the phenolic resin decomposes at 350°C.After infiltration with Si, the geometrical shape of the preforms was preserved and nofurther shrinkage was observed. Thus, the presented approach can be considered as a near-net-shape forming process of ceramic parts [19].XRD analysis after reactive infiltration revealed the presence of only two crystallinephases, Si and b -SiC. The as-infiltrated SiSiC laminar composites had a density of 2.3 – 0.02 g cm-³. According to the SEM micrographs presented in Fig.

The as-infiltrated SiSiC laminar composites had a density of 2.3 – 0.02 g cm-³. According to the SEM micrographs presented in Fig. 4 the porosity was com-pletely infiltrated with Si, however, a substantial amount of unreacted carbon still remained.The phase composition after infiltration as obtained from stereological analysis of the SEMmicrographs (Image C Intronic GmbH, Germany) was approximated to 18.3 vol% SiC, 75.4vol% Si and 6.3 vol% of residual carbon.a)b)SiC500 µm50 µ mSiCFigure 4: Microstructure of the SiSiC composite infiltrated 1h at 1500°C and annealed for 1h.262The Youngs modulus for the specimen infiltrated at 1500°C for 1h was determined as 165 –7 GPa. The measured values of the elastic modulus of the fabricated SiSiC composites are ingood agreement with the predicted values for the composite materials [20]:Ecomp∑=ivEiiwhere Ei is the elastic moduli of the constituents and Vi is the volume fractions.

Using thedate for the plain phases (ESi =167 MPa [20], ESiC = 414 MPA [21] and EGlassyC = 20 GPa[22]) a Youngs modulus of 200 GPa is estimated, which is slightly higher than the experimen-tal one due to a higher amount of residual carbon than obtain by image analysis.The infiltrated SiSiC composites obtained an average bending strength of 115 –15 MPa,which is in the same range as other SiC-material manufactured LOM of cast SiC-tapes[11,12]. Fig. 5 shows the correlation between the four point bending strength and the fiberorientation of the specimen. A fiber orientation of 0° represents a loading direction (tensilestress) parallel to the fiber elongation, whereas an orientation of 90° denotes a loading direc-tion perpendicular to the fiber. A fiber orientation of 0/90° was created by alternatively stack-ing layers with fiber orientation of 0° and 90°.

A fiber orientation of 0/90° was created by alternatively stack-ing layers with fiber orientation of 0° and 90°. The bending strength remained almost invari-ant with changing the fiber orientation which might be announced with the high amount of Siand the low fraction of SiC giving rise for a quasi isotropic mechanical behavior. Hence fur-ther development will be focused on reducing the fraction of Si and unreacted carbon.263140] a P M130[ h t g n e r t s g n d n e Bi12011010090800°90°0/90°LayerFiber orientationFigure 5: Bending strength vs. fiber orientation.Fig. 6 shows the rise in flexural strength with increasing SiC content. According to Fig 7which show the influence of the preform porosity on the phase composition of the component,the maximum SiC content is reached with a preform porosity of 42.7 vol%.

According to Fig 7which show the influence of the preform porosity on the phase composition of the component,the maximum SiC content is reached with a preform porosity of 42.7 vol%. The focus of fur-ther research is on decreasing the preform porosity by infiltration resins with high carbonyield or SiC-slurries into the as LOMed components, to increase the SiC- content and improvethe mechanical properties.450] a P M[400350ht g n e r t300S250la r u x e Fl200150100020406080100SiC fraction [vol%]Figure 6: Flexural strength vs. SiC content264s t c u d o r p1,00,8Chocking-off regionSiCSifo0,6n oi t c a r f0,4e m u o Vl0,20.4270,00,00,20,40,60,81,0Initial pore volume fraction of the carbon precursorFigure 7: Influence of preform porosity on the phase composition.ConclusionsSiSiC composites with a laminar structure were fabricated by Laminated Object Manufactur-ing.

Carbon sheets made from pitch fibers were coated with an adhesive consisting of pheno-lic resin and butyl nitrile rubber and processed in a LOM-device, subsequently pyrolysed andfinally infiltrated with silicon melt. Due to the extremely low dimensional changes, LOM ofcarbon sheets followed by post infiltration processing can be considered as a near net shapemanufacturing process. Reduction of residual Si and unreacted carbon is expected to give risefor improving the mechanical properties of the LOM SiSiC composite materials,AknowledgementsThe authors gratefully acknowledge the Deutsche Forschungsgemeinschaft DFG for theirfinancial support.265References[1] W.Krenkel, T.Henke, Design of high performance CMC brake discs, Key Eng. Mat.1999, 164, 421.[2] M.Wilhelm, M.Kornfeld, W.Wruss, Development of SiC-Si composites with finegrained microstructures, J. Eur. Ceram. soc.

vanced Process Planning for Subtractive Rapid Prototyping Joseph E. Petrzelka and Matthew C. Frank Department of Industrial and Manufacturing Systems Engineering The Iowa State University of Science and Technology Ames, Iowa, USAAbstract This paper presents process planning methods for Subtractive Rapid Prototyping, which deals with multiple setup operations and the related issues of stock material management. Subtractive Rapid Prototyping (SRP) borrows from additive rapid prototyping technologies by using 2½D layer based toolpath processing; however, it is limited by tool accessibility. To counter the accessibility problem, SRP systems (such as desktop milling machines) employ a rotary fourth axis to provide more complete surface coverage. However, layer-based removal processing from different rotary positions can be inefficient due to double-coverage of certain volumes.

However, layer-based removal processing from different rotary positions can be inefficient due to double-coverage of certain volumes. This paper presents a method that employs STL models of the in-process stock material generated from slices of the part along the rotation axis. The developed algorithms intend to improve the efficiency and reliability of these multiple layer-based removal steps for rapid manufacturing.Introduction Subtractive Rapid Prototyping (SRP) is a considerably lesser known and utilized form of rapid prototyping technology, mainly due to continued challenges in the pre-process engineering and setup planning required. Subtractive operations in general afford excellent accuracy and repeatability, which is why it is more often utilized as an additional function of a hybrid-type rapid prototyping system that includes both additive and subtractive operations.

Since the early days of RP technology, with systems such as Laminated Object Manufacturing (LOM), subtractive means have been employed as part of a solution alongside the conventional additive layer-based approaches. Although perhaps not purely subtractive, in LOM, a laser was used to cut the profile and surrounding support structures of each layer, after the layer of paper is added. In this manner, the additive approach enabled improved geometric capability while the laser cutting offered reasonable shaping accuracy. Later, the Sanders Modelmaker system, now used heavily in jewelry and dental manufacturing, utilized a machining process to accurately mill each deposited layer to a precise thickness. In a research project, Shape Deposition Manufacturing (SDM) used 5-axis machining in conjunction with a variety of deposition approaches to create both parts and molds.

In a research project, Shape Deposition Manufacturing (SDM) used 5-axis machining in conjunction with a variety of deposition approaches to create both parts and molds. More recently, systems such as LAMP and Ultrasonic Consolidation (UC) continue to use subtractive means in an iterative manner to improve accuracy and surface finish. Desktop milling machines suitable for rapid prototyping have been marketed for lower-end materials in single or multi- sided machining operations, while lower-end, but more user-friendly SRP software systems such as Millit and Deskproto have attempted to facilitate the NC programming for these applications. To date, subtractive processes continue to be utilized in a few RP systems, while SRP-only systems have had limited success. The overwhelming success of RP technologies to date can be attributed almost solely to the additive-only machines; however, there are niche applications where an SRP system would be desirable.

The overwhelming success of RP technologies to date can be attributed almost solely to the additive-only machines; however, there are niche applications where an SRP system would be desirable. This paper presents one of several new methods the authors have developed for SRP process planning, specifically one that will enable more efficient execution and reduced processing times.Subtractive processes in general are extremely versatile and can form a variety of complex geometries from a vast array of materials. However, to form arbitrary geometry from stock, material must be removed from a variety of orientations, or setups. For example, if one were to mill a part from a block of stock material, several setups would be required; one possibility, though not necessarily optimal, would be to mill the part from orthogonal angles (Figure 1). In situations such as this, it becomes very difficult to analyze the remaining stock after a certain sequence of operations.

In situations such as this, it becomes very difficult to analyze the remaining stock after a certain sequence of operations. The problem is far from trivial; component geometry will often occlude accessibility to certain portions of the stock material. A varietyof tool path simulation and verification techniques have been presented in the literature; however, these techniques focus computational on effort surface finish and conformity than rather shape gross analysis.(a)(b)(c)(d)Figure 1 – Multiple setups required (even in multi-axis milling) to create component via a subtractive process; (a) shows initial stock geometry and (b-d) show the results after iterative cutting operations until (d) the component fully emergesLittle research has been conducted that specifically evaluates the effect of multiple-setup roughing operations; in fact, many typical situations do not require this type of analysis. In usual production settings, a near-net shape component is created (i.e.

In usual production settings, a near-net shape component is created (i.e. casting), and then critical features are machined. In die and mold shops, a single setup in a multi-axis mill is typically sufficient to achieve the geometry of the die or mold cavity. Regardless, there exist a variety of applications where multiple setup subtractive analysis is a requirement, especially for creating custom or unique parts from nominally sized stock material. In this research, we consider a subset of this problem where all setups can be performed by rotation about a single axis.

In this research, we consider a subset of this problem where all setups can be performed by rotation about a single axis. The manner of iteration is not important; the work piece could be fixtured in a four-axis machine and automatically rotated to new orientations, or the work piece could be manually fixtured in a three-axis mill such that each subsequent orientation was about a theoretical axis (as in Figure 1).(a)Recent development of a rapid prototyping process using CNC machining by the authors [8] has enabled automated process planning for custom, very low volume, or prototype components. In this system, support structures provide fixturing for machining in a four- axis milling machine. Visibility analysis of cross sectional slice data provides a basis for automated setup planning about a single axis.

Visibility analysis of cross sectional slice data provides a basis for automated setup planning about a single axis. The implementation uses a modified Greedy set cover algorithm to determine discrete orientations; depth requirements are generated by ‘bucketing’ visible slice segments for each orientation. In its full implementation, CNC-RP safely and completely machines arbitrary components from bar stock in a four-axis mill. In effect, the fourth axis serves as a rotary setups between to automate independent three-axis tool paths (Figure2).called CNC-RP,sacrificialindexerRotary IndexerEnd MillRound StockAxis of RotationOpposing 3-Jaw Chucks(b)(Side View) (Side View)(1)(5)(2)(6)While this system has proven capable and robust, it generates each tool path independently of others, basing it only on set cover angles and depth results. The conservative tool path planning method results in efficiency sacrifices for this multiple orientation process.

The conservative tool path planning method results in efficiency sacrifices for this multiple orientation process. Because each machining orientation is performed independent of others, there is no knowledge of prior operations during roughing(3)(4)Final Steel Part(b) Figure 2 Process sequence of steps (b.1-b.4) to expose component geometry and (b.5-b.6) to expose sacrificial supports(a) CNC-RPImplementation;procedures. This naïve approach leads to excessively long tool paths and redundant volume sweeps (Figure 3). If an effective means existed to predict the stock geometry resulting from each step (setup) in the process, one could generate much more effective and efficient tool paths.

If an effective means existed to predict the stock geometry resulting from each step (setup) in the process, one could generate much more effective and efficient tool paths. Saito [18] cites machining time as one of five critical to successfully producing complex geometries; naïve tool path generation fails to minimize this factor.factors(a)(b)Figure 3 - (a) Naïve tool path volume sweep; (b) Volume sweep of tool path with knowledge of prior operationsTo characterize the tradeoff between robust (naïve) and fast characteristics, we define two types of error that may occur when generating the in-process stock material shape; Type A: Remaining stock is underestimated, where the failure mode is possible collision between tool and large amount of stock, and, Type B: Remaining stock is overestimated, where the failure mode is a marginal increase in tool path length and machining time.

Noting the failure modes and the desire to develop a robust system, it is asserted that Type A error must be avoided at all costs. For example, in the lights-out operation of a Subtractive Rapid Prototyping method like CNC-RP, it is absolutely necessary that tool crashes do not occur during untended operations, often done overnight. Without simulation of previous operations (three axis tool paths at different orientations), each subsequent tool path is required to assume worst-case conditions (Type B error). While little original stock material may actually remain, a large volume is swept by the tool to maintain a robust process (Figure 3a). Commercially available CAM packages can accept STL (triangle mesh surface) models as stock material input to constrain tool path generation to the correct regions in space. Attempts to utilize existing simulation techniques in CAM packages to generate this driving STL model have not been successful.

Attempts to utilize existing simulation techniques in CAM packages to generate this driving STL model have not been successful. Z-buffer based simulations provide reasonably accurate results, though not robust. After several iterations between the Z-buffer dexel structure to a portable STL format (for use simulating several sequential machining operations), the data becomes unreasonably large and prone to corrupt STL files. Furthermore, this and similar simulation techniques rely on pre-generated tool paths.

Furthermore, this and similar simulation techniques rely on pre-generated tool paths. These difficulties create the need for a simple remaining stock calculation for multiple setup environments that is (1) mathematically defined, with portable results, (2) computationally inexpensive for application to many unique components, and (3) able to predict process performance prior to computationally expensive tool path planning.Related Work Algorithms for subtractive process planning can be classified as (1) tool path generation methods or (2) simulation / verification methods [18]. The literature fails to address the area of ‘lightweight’ pre-NC computation (gross shape analysis only, neglecting surface scallops) for process planning in a multiple- setup environment. Tool path simulation is similar to the remaining stock calculation desired; it creates a model of the stock by subtracting the tool’s swept volume from a specific tool path.

Tool path simulation is similar to the remaining stock calculation desired; it creates a model of the stock by subtracting the tool’s swept volume from a specific tool path. Verification techniques go a step beyond to detect gouges (over-cutting during feed motion), under-cutting (failing to remove all material from a component surface), and collisions (driving to tool into a part or fixture during rapid travel). By definition, simulation and verification calculate (in various manners) the Boolean effect of a tool path. Three primary groups of verification / simulation are evident in the literature: vector based, spatial discretization, and solid/parametric surface based methods. Chappel [5] and Oliver/Goodman [16] developed vector based approaches to track machining progress. After populating component surfaces with normal vectors (the density determined relative accuracy and computation time, though in O(n)), the vectors are trimmed as appropriate for each tool instance.

After populating component surfaces with normal vectors (the density determined relative accuracy and computation time, though in O(n)), the vectors are trimmed as appropriate for each tool instance. Jerard [12] extended this simulation technique to verification and cutter location (CL) data correction. The most popular spatial discretization method – the Z-buffer - was originally proposed by Van Hook [19]. Huang and Oliver [11] extended this simulation technique to verification by detecting gouging and undercutting. Saito and Takahashi [18] presented a similar graphics-based approach (G-buffer method) that is capable of both tool pathgeneration and simulation. Bohez et al. [3] discretize space into x-y planes in their method. Solid modeling would seem to be ideal for NC simulation and verification; an exact representation of remaining material could be constructed.

[3] discretize space into x-y planes in their method. Solid modeling would seem to be ideal for NC simulation and verification; an exact representation of remaining material could be constructed. However, constructive solid geometry (CSG) application becomes O(n4) and therefore unacceptable for any simulation of practical scale [12]. There does not appear to be extensive recent research in this area; most efforts have been focused on parametric boundary representation (Brep) methods (effectively a transition from solid to surface modeling). Using Brep methods, Weinart [21] achieves 5-axis swept volume approximation resulting in parametric NURBS surfaces. Boolean subtraction of this swept volume from a stock model results in a tool path simulation. Weinart still interpolates between discrete tool instances (similar to discretization methods) and shows only a small example with no large scale implementation.

Weinart still interpolates between discrete tool instances (similar to discretization methods) and shows only a small example with no large scale implementation. Jütler [13] presents a similar method, though he used discrete points along a rational B-spline curve to drive tool instances. Jütler presents purely empirical research with no successful implementation. Blackmore [2] generates triangle mesh representations of swept volume using differential equations applied to continuous tool paths and boundaries. All of these Brep methods are slow [3] and do not appear to be mature technologies based on the lack of full scale implementations in the literature. Lauwers [15] and Huang [11] both use simulation results to correct CL data (tool position & posture). Pal [17] presents a remaining stock computation to drive more efficient tool paths, but his algorithm is designed only for two-stage three axis milling with sequentially smaller ball mills.

Pal [17] presents a remaining stock computation to drive more efficient tool paths, but his algorithm is designed only for two-stage three axis milling with sequentially smaller ball mills. Many of these simulation techniques are well suited to verify gouging and undercutting in surface finishing operations. However, most are very slow (Z-buffer methods are the only that can be considered fast [3]) and all rely on computational investment in tool path generation. The literature does not present methods suitable for gross stock remainder calculations for more efficient roughing processes. Only vector based approaches offer mathematically defined tool tracking with respect to the surface (though some other methods can derive it from discretized maps).Methodology In this section, we provide a general description of the proposed methods for remaining stock calculation.

The approach involves three major tasks; (1) the model of the stock material is broken into cross sectional slices and a set of operations performed to create a factor of safety for avoiding Type A errors, (2) slices are modified to simulate the iterative changes to the stock, and finally (3) a reconstruction of the slice data into a surface model representation of the stock for each iterative step in the subtractive rapid prototyping process. The method presented computes the remaining stock for a set of subtractive processes where all setups occur about a single theoretical axis. In particular, this analysis computes the net effect of a tool path from a given orientation to a given depth, but without requiring exact tool path data. If a robust tool path generator is used, the only gain from considering tool instances is the ability to analyze surface scallops; for a roughing operation this is considerably less important than in finishing operations.

Our method consists of three distinct algorithms: slice approximation, slice shadowing, and polyhedral reconstruction (Figure ).∩=Prototype ModelSlice ApproximationSlice ShadowingPolyhedral ReconstructionFigure 4 – Method for remaining stock computation, composed of three stepsSlice Approximation Slice approximation is a pre-processing step used to convert a surface model of the part geometry to cross sectional polygonal data. Since we assume all setups occur about a single theoretical axis, the problem can be represented as a set of cross sections perpendicular toand along that axis. This provides the distinct advantage of reducing the overall problem to independent analyses of two-dimensional data from a three dimensional model. However, the preprocessing involves more than simply slicing the model.

However, the preprocessing involves more than simply slicing the model. Subsequent algorithms in the method use a visibility based shadow to approximate remaining stock, so the slices must account for the difference between visible and accessible regions. If a visibility approach is used on the exact cross section of the part, Type A error will occur for all but the simplest geometries (remaining stock will be underestimated). To avoid this, the cross sectional data must exclude small concave features (i.e. holes and slots) that a tool of specific finite diameter may not be able to access. The approach to counter visibility underestimation is twofold: cross sections are (1) reduced to their own convex hull polygons and (2) redefined as the union of all slices within one tool radius along the axis of rotation. The first measure has the effect of closing concave geometry that is radial relative to the axis of rotation, while the second measure closes concave geometry along the axis.

The first measure has the effect of closing concave geometry that is radial relative to the axis of rotation, while the second measure closes concave geometry along the axis. As the final step of slice approximation, the resulting cross sections can be offset in order to approximately account for scallop effects. Since specific tool instances are assumed unknown, the location of scallops is likewise unknown. To this point, offsetting the polygonal cross sections allows compensation for possible scallops and for any amount of intentional under-cutting (e.g. leaving some set amount of material for finishing operations). The net effect of the slice approximation steps described above is shown in Figure 5 with a reconstructed model. In particular, note slots inaccessible by a roughing tool have been eliminated; however, thick features have been slightly over approximated (Type B error).

In particular, note slots inaccessible by a roughing tool have been eliminated; however, thick features have been slightly over approximated (Type B error). This reduction in accuracy is the cost of using a robust form of visibility analysis during slice alteration; however, it provides a critical degree of safety to ensure tool collisions will not occur.thatthesmall holes and(a)(b)Figure 5 – Example component along axis of rotation, (a) original model and (b) reconstruction of approximating slices showing elimination of inaccessible holes and slotsThe methodology begins with an STL model of a part and fixturing geometry (representing the space set in which no machining should take place) is sliced into sets of contours using existing slicing methods. For multiple setup processes, the cross sectional slice planes are set normal to the axis of rotation.

For multiple setup processes, the cross sectional slice planes are set normal to the axis of rotation. For this work, we will define of n slices, each denoted Ci, where l is the total length of component along rotational axis, d is the distance between slices and r is the tool radius. Since this methodology is later used to drive tool path planning, it must never have Type A error so that collisions do not occur; equivalently, the amount of stock must never be underestimated. To maintain a lightweight model and avoid the computational complexity associated with collision or machinability algorithms, two steps are taken the slices and avoid underestimating material removal in visible regions that a tool could not actually reach. First, each cross section its own convex hull (Ci=ConvexHull(Ci) ; i=1…n) as shown in Figure 6.

First, each cross section its own convex hull (Ci=ConvexHull(Ci) ; i=1…n) as shown in Figure 6. This step has the effect of eliminating slots or other internal geometry radial with respect to the axis of rotation (any given convex hull is completely visible and accessible by any convex 2D tool silhouette). Second, each slice is redefined as the union of itself and all slices within one tool radii; (Ci=(Ci-(r/d)) U (…) U (Ci) U (…) U (Ci+(r/d)) ; i=1…n), as shown in Figure 7.: {C1, C2, … , Cn}as the component data setremaining(a)(b)to further processFigure 6 – Each slice (a) is represented as (b) its convex hull polygonisreducedto(b)(a)Figure 7 – Each slice in (a) is redefined as the union of slices within one tool radius of the slice (as shown in image b for the bold slice in image a)This second step offsets all possible collision surfaces by one tool radius and closes small gaps and slots along the axis of rotation.

Both of these steps contribute Type B error, where the remaining stock is over- estimated, though not detrimentally so as discussed in the implementation section. In addition to the part geometry, slices must also be created to represent initial stock geometry. Slices can be generated automatically to represent homogenous stock material (bar or prismatic stock), or a model of the beginning stock can be sliced to create cross-sections. Regardless, the beginning stock is assumed to be of convex cross section along the axis of rotation, where we define: : {S1, S2, … , Sn}, the stock data set of n slices, each denoted Si. It should be noted that for stock material that is homogenous along the axis are equivalent.of rotation (i.e. bar or round stock), all Si inSlice Shadowing Slice shadowing, the second of three steps, is central to the remaining stock analysis.

bar or round stock), all Si inSlice Shadowing Slice shadowing, the second of three steps, is central to the remaining stock analysis. After the model has been represented as a set of cross sections, each cross section is independently modified with a unique ‘limited visibility shadow’. Unlike infinite visibility from a given orientation, we explicitly define a finite visibility range to emulate the finite cutting depth common to subtractive processes such as machining. This shadow is readily defined by Boolean operations between the part cross sections and initial stock cross sections. The algorithm is generalized to allow multiple ‘cuts’ to different depths from different orientations.Once the component has been approximated as a series of slices, , the effect of cutting is simulated on each slice Ci independently.

For this simulation, we note that the remaining stock will be equivalent to the union of the stock material (cut to a certain depth), the component, and the visibility shadow of the component (where the tool could not reach). Recall that measures were taken during ‘Slice Approximation’ to account for the difference between visibility and accessibility analysis.

Recall that measures were taken during ‘Slice Approximation’ to account for the difference between visibility and accessibility analysis. For the purposes of detailed discussion, we further define m as the number of cutting orientations, : {α1 , α2 , … ,αm}, the set of cutting angles, : { δ1 , δ2 , … , δm} , the set of cutting depths, relative to axis of rotation viewed from angles j : {A1,j , A2,j , … , An,j}, the set of n intermediate slices, representing the visibility shadow from angle αj , j : {B1,j , B2,j , … , Bn,j}, the set of n intermediate slices, representing stock material (without component or shadow) from angle αj to depth δj, j : {R1,j , R2,j , … , Rn,j} , the set of n slices representing effect of a single cut from angle αj to depth δj and finally j : {ReSt1,j , ReSt2,j , … , ReStn,j}, the set of n slices, representing effect of cumulative cuts from 1 … j. With the final goal of achieving slice setj , setsj ,j , and j are computed.

With the final goal of achieving slice setj , setsj ,j , and j are computed. Each is presented below:Slice set j (Figure 8) represents a visibility shadow from angle αj cast onto set . Because each Ci is a known polygon, a visibility shadow can be easily determined by finding points in Ci tangent to αj.. To find these points, some vector in the direction of αj is determined. Then, the cross product between this vector and each segment (span between polygon vertices pk and pk+1) in the cross section is found; if the sign of this cross product changes between two segments, their shared endpoint is tangent to αj. In the case presented, the polygonal cross section is known to be convex, thus, there are exactly two points tangent to any given direction vector. Two new vertices, pu,j and pv,j, can be added to the polygon point set at some arbitrarily large distance in the direction of αj from tangency points pu and pv).

Two new vertices, pu,j and pv,j, can be added to the polygon point set at some arbitrarily large distance in the direction of αj from tangency points pu and pv). Treating the polygon vertices as a point cloud and recomputing the convex hull yields a ‘shadow’ contour to some finite distance (the general case of non-convex polygons would require a simple modification to detect the intersection of the shadow ray and other polygon segments). Finally, this ‘shadow’ is trimmed to the stock cross section via intersection and the result is stored as Ai,j.SiShadow (Ci)jAi,jj∩=Figure 8 - shadow of Ci at angle αj trimmed to stock cross-section Sij is the set of slices Ai,j representing a visibilityNext, slice set j (Figure 9) represents the effect on stock material of a cut from angle αj to depth δj, assuming for now that tool no component geometry restricts access (we are only considering stock cutting for j).

First, an arbitrarily large polygon is created to represent the half- space not accessible by a tool at depth δj at angle αj. In practice, this is done by forming a square sufficiently centered on the origin, translating this square polygon so that the top segment is at depth δj, and rotating the polygon by angle αj about the origin. Then, each Bi,j can be computed by trimming, via intersection, the stock cross section (Si) to the appropriate ‘half-space’ polygon. In the special case that the stock is uniform along the slicing (cross sectional) axis, all Bi,j are equivalent for each cut j and only one Bi,j need be computed.‘Halfspace’jBi,jSij∩=j is the set of slices Bi,j composed of stock material Figure 9 - unreachable from angle αj to depth δj , computed by trimming the unreachable ‘halfspace’ to the stock cross-sectionlargej yields j,(Figure 10) representing the net effect of a single cut from angle αj to depth δj.

Without formal proof, we assert that this represents the remaining stock assuming that a tool path is used that has no gouge (removing too much material) or undercut (failing to remove all desired material) motion.The union of respective slices inj andAi,jBi,jRi,jjU=Figure 10 – Single-cut effect is represented by slice j, where each Ri,j is computed by joining set shadow Ai,j and stock material Bi,jFinally, to evaluate the cumulative effect of cutting operations, the set sense, one would use each However, this makes computations dependent on each other; it would be more appropriate to compute each 1 … j can be computed independently for large scale simulations. Thus, the cumulative effect of cuts ( j) , (Figure 11) is computed as the cumulative intersection of single-cut effects ( j), allowing independent calculation of single-cut effects regardless of operation order.j is computed.

In the natural j as the beginning stock material for computing the effect of cut j+1.Ri,jRi,j+1ReSti,j+1j in terms of the precedingj such that all∩=Figure 11 – Each ReSti,j in cumulative effects Ri,j, allowing independent computation of Ri,jj+1 is computed as the intersection of preceding single-cutPolyhedral Reconstruction The resulting contours from the shadowing operation compose a rich data set, though not portable in any widely accepted format. Hence, the third and final step is a polyhedral reconstruction of the stock geometry from the modified slice data, which converts the data to a common format (STL model) for graphics rendering and use in CAD/CAM packages. Polyhedral reconstruction from polygonal cross sections is a relatively well known problem, and in this work we created our own methods unique to this application. Our approach addressed the three primary problems of reconstruction: correspondence, branching, and tiling [1].

Our approach addressed the three primary problems of reconstruction: correspondence, branching, and tiling [1]. However, for brevity we will not include a description of these methods. The primary focus of this paper was the remaining stock calculations, enabled by slice approximation and slice shadowing methods described above.j+1Implementation The algorithms presented above have been implemented in software and utilized in the laboratory for four-axis subtractive rapid prototyping (CNC-RP). These algorithms are ideally suited for this system because they provide a robust, reliable estimation of remaining stock. In the CNC-RP process, it is more important to achieve safe and effective results for first-pass success in low volume applications than it is to optimize tool paths at a high level (the expectation is that only one, or very few parts are desired).

This remaining stock calculation serves exactly that purpose: while the stock is overestimated in certain situations, it is done so to ensure the integrity of results and eliminate collisions between the tool and underestimated stock material while the machine is allowed to run unattended. To illustrate the results typical in the CNC-RP implementation, reconstructed models for a single component machined in seven setups are illustrated in Figure 1.Prototype ModelRound Stock…after ops 1-4+ operation 5+ operation 6+ operation 7Figure 12 – Remaining stock approximations for the example component from seven setup orientationsThe efficiency gains from this implementation have been very significant. While a variety of factors may influence tool path inefficiency, one of the largest contributors is adopting a naïve and overly conservative remaining stock assumption.

While a variety of factors may influence tool path inefficiency, one of the largest contributors is adopting a naïve and overly conservative remaining stock assumption. Overall tool path length has been observed to decrease as much as 65% when utilizing this new remaining stock analysis during tool path generation.Improved efficiencies are presented for two sample parts (Table 1). Each was assumed to be machined from three inch diameter bar stock with sacrificial supports (consistent with the CNC-RP implementation) using a 0.5 inch flat end mill, 0.04 inch step down, and 75% step over. In laboratory testing, the tool was previously observed performing redundant ‘air’ cutting, but with the remaining stock implementation it remains engaged in the material during nearly the entire roughing routine.

Numerous laboratory tests on a variety of geometries have not revealed any cases of Type A error (failure mode of collision with under- estimated remaining stock).Table 1 – Results (coupling with multi-axis tool path generation) for two example geometriesCAD ModelReconstructed Remaining Stock ModelTotal SetupsTool Path Length ReductionComputation Time (3 GHz CPU)953%640 milliseconds864%640 millisecondsLimitations and Future Work This work is limited primarily by the assumption that setups must be about a single axis. However, in practice, it is observed that most geometries can be achieved in this manner (from machinery components to free-form biologic shapes such as human bones). This fact may be due to the motivation behind design for manufacturing to reduce setups for industrial components. Using visibility software in the lab, the authors have found few bones of the human anatomy (i.e.

Using visibility software in the lab, the authors have found few bones of the human anatomy (i.e. the entire skull and pelvis) that cannot be manufacturing about one axis of rotation. The steps taken during slice preprocessing to avoid underestimating the remaining material (approximation as a convex hull) cause slight overestimation in many cases. This constraint could be relaxed if a machinability algorithm was incorporated and a more robust correspondence algorithm implemented for polyhedral reconstruction. While this would result in a more exact approximation, the machinability algorithm would substantially increase computation time. As a final limitation, the method presented does not convey scallop effects or surface finish in the remaining stock estimation.

As a final limitation, the method presented does not convey scallop effects or surface finish in the remaining stock estimation. However, the method presented is not intended to replace conventional verification and simulation techniques for finishing operations; rather, its strength is in advanced, efficient process planning.Jerard [12] proposes a framework for automated NC code generation via an iterative process between setup orientations, setup sequencing, tool path generation, and tool path verification. With the algorithms we present, setup orientation and sequencing can be improved without iteration between computationally expensive tool path generation and verification algorithms. The robust and simple nature of this algorithm lends itself to future simulation or optimization techniques for determining optimal setup strategies.

The robust and simple nature of this algorithm lends itself to future simulation or optimization techniques for determining optimal setup strategies. The Z-buffer method was the fastest pre-existing simulation method [3]; Huang and Oliver [11] cite a computation speed of 68.8 instances / second for a three-axis roughing routine (analogous to the operations analyzed here). Even at this rate, the tool path for the example part would have taken several minutes to simulate (for 9069 linear inches of feed move in the example component’s NC program), even at modern computing speeds. However, by evaluating the net effect of the tool path rather than each individual move, our method requires only 0.6 seconds to simulate nine different operations from different setup orientations (considering the example part on a 3 GHz CPU).

Because the net effect of a tool path requires only milliseconds to calculate, this algorithm is appropriate for application to optimization techniques.One such application for focus of future work is for the detection of thin material conditions. If operations are not sequenced optimally, machining from opposing setup angles (at or near 180°) can form thin webs or thin strings (Figure 13). These conditions can be readily detected by our algorithm at the slice level.thin stringthin webFigure 13 – Examples of thin webs and strings emerging from the multi sided machining of a human femur from round stockThe method presented can also be applied to design analysis. The shadow method can be used to augment visibility algorithms. While visibility and accessibility algorithms investigate the invisible / inaccessible surface area of a component, our new method can be used to evaluate inaccessible volume from a given set of angles.

One possible application would be evaluating the amount of material that must be added to regions of a component to make molding or forging possible from a minimal number of molds. Another future application is to simulate wire EDM processes since, in fact, wire EDM cannot readily create concave geometry and it is very nearly a line of sight accessible process (for small wire diameters). Thus, the assumptions made for the current algorithm would map directly to a wire EDM process.Conclusions This work presented a new method for gross analysis of subtractive processes in multiple setup environments. As opposed to existing methods, focusing on single setup (albeit multi-axis) surface finishing verification, our method provides a robust and efficient means of evaluating subtractive processes. Specific results of interest are a new limited visibility shadowing algorithm and efficient analysis of subtractive processes independent of specific tool paths.

Specific results of interest are a new limited visibility shadowing algorithm and efficient analysis of subtractive processes independent of specific tool paths. The limited visibility shadowing algorithm can be applied to a variety of other situations. Other visibility analyses evaluate surface visibility, and a set cover problem can be formulated to solve for optimal setup angles. Using this algorithm, the problem could be formulated differently to achieve the maximum accessible volume in a set cover problem rather than maximum accessible surface area. Moreover, the extremely fast nature of this algorithm lends itself to iterative optimization approaches; which has significant implications to improved process planning.

Moreover, the extremely fast nature of this algorithm lends itself to iterative optimization approaches; which has significant implications to improved process planning. If an objective function can be formulated that uses metrics to identify advantageous or detrimental characteristics of process progress, the millisecond time scale of our algorithms will allow many iterations through setups to identify (at least locally) optimal setup sequences. This could be specifically advantageous in 4-axis SRP, if the desired prototype has complex geometry that will require numerous arbitrary setup orientations about the axis of rotation.Acknowledgment Financial support was provided by grants from the National Institutes of Health (AR48939 and AR55533) and Deere and Company (Account 400-60-41).References[1][2]Bajaj, C., E. Coyle, and K.-N. Lin. “Arbitrary Topology Shape Reconstruction from Planar Cross Sections”, Graphical Models and Image Processing, v58 n6 November 1996.

Coyle, and K.-N. Lin. “Arbitrary Topology Shape Reconstruction from Planar Cross Sections”, Graphical Models and Image Processing, v58 n6 November 1996. Blackmore, D., M. C. Leu, and L. P. Wang. “The sweep-envelope differential equation algorithm and its application to NC machining verification”, Computer-Aided Design, v29 n9 1997.[3][4][5][6][7][8][9]Bohez, E., N. Minh, B. Kiatsrithanakorn, P. Natasukon, H. Ruei-Yun, L. Son. “The stencil buffer sweep plane algorithm for 5-axis CNC tool path verification”, Computer-Aided Design, v35, n12, October 2003. Boissonnat, J.-D. “Shape Reconstruction from planar cross-sections”, Computer Vision, Graphics, and Image Processing, v44 n1 October 1988. Chappel, L. “The use of vector to simulate material removed by numerically controlled milling”, Computer-Aided Design, v15 n3 1983. Christiansen, H. and T. Sederberg. “Conversion of Complex Contour Line Definitions into Polygonal Element Mosaics”, Computer Graphics, v12 n3 1978.

and E. Goodman. “Color graphic verification of N/C milling programs for sculptured[17][18]surface parts”, First Symposium on integrated intelligent manufacturing, New York, ASME 1986. Pal, P. “Remaining Stock Computation for 3D-Machining in Parametric Regime”, Journal of Manufacturing Science and Engineering, v127 n4 November 2005. Saito, T. and T. Takahashi. “NC Machining with G-buffer Method”, Computer Graphics, v25, n4, July 1991.[19] Van Hook, T. “Real-Time Shaded NC Milling Display”, Computer Graphics, v20 n4 August1986.[20] Wang, Y. and J. Aggarwal. “Surface Reconstruction and Representation of 3-D Scenes”, PatternRecognition, v19 n3 1986.[21] Weinert, K., S. Du, P. Damm, M. Stautner. “Swept volume generation for the simulation of machining processes”, International Journal of Machine Tools & Manufacture, v44 n6 May 2004..

VANCED SHEET METAL MANUFACTURING USING RAPID TOOLINGDu Zhao Hui1, Chua Chee Kai1, Chua Yew Sen1 Loh-Lee Keow Gek2, Lim Ser Tiak2 1 School of Mechanical & Production Engineering, Nanyang Technological University, 50 Nanyang Avenue, Singapore 639798 2School of Engineering, Temasek Polytechnic, 21 Tampines Ave 1, Singapore 529757ABSTRACTA closed loop process is proposed for making sheet metal prototyping parts by using advanced computer aided techniques and computer controlled machines. The key aspect of this process is the method used to fabricate and modify the sheet metal forming tools, which are not necessarily for mass production but should be suitable for short run production or design evaluation of sheet metal products where the prototyping cost and lead-time are greatly reduced.Various approaches are investigated in the preparation of the tooling for onward embossing on a sheet metal.

The three indirect approaches use Selective Laser Sintering (SLS), Stereolithography(SLA), and high speed Computer Numerical Controlled (CNC) milling to build the masters from computer data models. And the masters are then served in the vacuum casting process to generate the non-ferrous tooling. The direct approach uses DTM’s RapidSteel to produce the metal tooling without going through any secondary process. Comparisons on quality, leading time and cost are presented.Keywords: Sheet metal manufacturing; Rapid prototyping; Rapid tooling;RapidSteel1. INTRODUCTIONParts from rapid prototyping (RP) / rapid tooling (RT) [1,2,3] could be produced not only for engineering prototyping, but also for marketing purposes by distributors and dealers. With RP and RT, plastic injection molds have been created to produce prototyping parts by many academic institutes and industrial organizations in a bid to reduce leading time and cost.

With RP and RT, plastic injection molds have been created to produce prototyping parts by many academic institutes and industrial organizations in a bid to reduce leading time and cost. However, only a few developmental results have been reported so far on sheet metal products by using RP and RT [4,5]. Berger, et al. [6] discussed that the process of sheet metal forming can be improved with the integration of concurrent engineering methods and RP technology, and one case study was presented using easily castable amorphous material and integrating Stereolithography in the fabrication of the casting shell. However, the more result and information desired are still unavailable.522It is well known that the tooling preparation is the most crucial stage for a successful sheet metal manufacturing.

However, the more result and information desired are still unavailable.522It is well known that the tooling preparation is the most crucial stage for a successful sheet metal manufacturing. Conventional methods for manufacturing a forming tool usually adopt a trial and error-prone approach, therefore hardly to obtain the correct tooling shape and forming conditions probably because of spring back in the deformed sheet metal, material non-linearities etc. Often few calculation or simulation of forming operations can be done because of complex mathematical constraints and unpredictable material behaviours. If a part shape variation or a specific manufacturing requirement occurs, the whole process must be reiterated and the forming dies be rebuilt. The die building process often requires Numerical Controlled (NC) programming which is time consuming, especially when the design has a large number of complex surfaces.

The die building process often requires Numerical Controlled (NC) programming which is time consuming, especially when the design has a large number of complex surfaces. Therefore, the forming dies for small-batch series and prototypes cause disproportionally high costs or risks [7].There exists a strong demand in industries for rapid tooling of sheet metal forming with a shorter time and lower cost. Eventually, the high risk of unsuccessful design and tooling becomes avoidable. The tooling might not be strictly desired to produce a large quantity of products like the existing one does, but it must be of good service for a short-run production. The prototyping parts produced by this means serve as a functional tool in formability study, design iteration, process optimization, or even as final products for customized demands. Therefore, the concept of Rapid Sheet Metal Prototyping (RSMP) emerges.2.

Therefore, the concept of Rapid Sheet Metal Prototyping (RSMP) emerges.2. RAPID SHEET METAL PROTOTYPING (RSMP)Present RP techniques are not intended for forming of sheet metal products directly but they are useful in transforming the CAD model into the final sheet metal product with additional steps. The RSMP includes certain additional methodology to complete the transformation to sheet metal parts with the aids of advanced computer aided techniques and computer controlled machines. There are three major steps involved in the closed loop system (shown as in Figure 1).523Figure 1. The closed loop of the RSMP systemFor a start, a conceptual product or tool is modelled by using computation and analysis function in one three-dimensional (3D) CAD software. This is an essential step that is carried out at the outset to ensure the viability of the project.

This is an essential step that is carried out at the outset to ensure the viability of the project. Physical Domain CreationVirtual Domain CreationOnce the 3D model is developed, techniques such as RP and RT, high speed milling machine etc., serve for creating the soft tooling, which will in turn be functioned in a short run of sheet metal forming. However, the tooling is not hard enough to be employed in a mass production. Functional Stamped PartsAfter functional parts, or prototyping parts, are produced, they may be fitted into an assembly or analysed for quality and appearance. At this stage, tool design and embossed parts can be evaluated. The geometry of the functional parts can be reconstructed with the measurement data by either contact or non-contact 3D scanners. By importing the reconstructed model into the virtual domain, a comparison can then be made between the design model and actual product.

By importing the reconstructed model into the virtual domain, a comparison can then be made between the design model and actual product. Features such as the spring-back for bending and least die radius curvature or other errors will be figured out in CAD systems. By using this feedback loop and computer-aided verification (CAV), modifications can be made to compensate and eliminate errors in the next design iteration of tooling and functional parts. Currently, much attention is concentrated on finding alternative ways to create the special tooling. Consequently, two major approaches are undertaken to explore the methodology with the aids of RP. The indirect methods involve several conversion steps before the required tools are prepared. The masters have to be made before going through casting to produce the aluminum/epoxy tooling.

The indirect methods involve several conversion steps before the required tools are prepared. The masters have to be made before going through casting to produce the aluminum/epoxy tooling. Here, masters refer to models that posses the physical outlook of the desired parts but lack of the material properties such as tensile strength, hardness, durability, etc. In this case, three processes selected to make the masters are Stereolithography (SLA), Selective Laser Sintering (SLS), and high speed CNC milling.

In this case, three processes selected to make the masters are Stereolithography (SLA), Selective Laser Sintering (SLS), and high speed CNC milling. The direct approach for rapid tooling is implemented by using the DTM's metal sintering in order to avoid any intermediate step.The various approaches undertaken can be seen as the four paths A, B, Cand D in Figure 2.524Cencept designDetail designTool designApproach AApproach BApproach CApproach DPrototyping by SLSPrototyping by SLAPrototyping by high speed CNC millingPrototyping by DTM's RapidsteelSilicon rubber mouldingEP250 castingMachining/ polishingAssemblyEmbossingFigure 2. The flow diagram of the four approaches3. DESIGN CONSIDERATIONS3.1 Embossed DesignA Merlion logo (a lion head with a fish body/tail) is selected as the embossed design for all the approaches. The 2D design is first generated using Pro-Engineer Sketch module.

The 2D design is first generated using Pro-Engineer Sketch module. The text is next incorporated into the design such that it carries the meaning of NTU (Nanyang Technological University) and TP (Temasek Polytechnic) where both of them share the same lettering 'T'. This is shown in Figure 3. A good design should avoid the possibility of crumpling the sheet metal and disfiguring the shape of the merlion head and the lettering.525Figure 3. The embossed design3.2 Tooling DesignThe Turret Punch Press driven by servo-motors will be used for the embossing process. By modification of the standard forming tool to incorporate the concept of die insert, various embossed designs could later be changed easily. Once this concept is established, the die insert and punch are modeled using Pro-Engineer. After the modeling, the top surface of the die insert and punch is then ready for placement of the embossed design.

After the modeling, the top surface of the die insert and punch is then ready for placement of the embossed design. Each forming surface on the die insert is then subsequently offset by half the material thickness followed by a mirror function to get an opposite effect of the embossed design. The radius curvature of corners is initiallyset to be 0.6mm which can be reiterated later in the process to obtain a good value for the embossing operation.Figure 4. The design of the die insert526One important point for the tooling design is the alignment between the die insert and the punch. Slight misalignment will affect the quality of the embossed design and tool fracture will result in serious misalignment. Alignment problem in this case is solved by incorporating pin holes on the punch and die insert. A die holder for holding the die insert and a stripper are next designed in Pro-Engineer to fit that of the new concept.

A die holder for holding the die insert and a stripper are next designed in Pro-Engineer to fit that of the new concept. Since both the die holder and stripper, the exchangeable parts in the whole die, only need to be produced once, they were produced using conventional machine instead of RT technology by an external contractor.In order to ensure viability of the design, the forming tools with the holder and stripper are assembled in Pro-Engineer to verify that they fit nicely with each other and virtually simulate the embossing process. The design of the die insert is presented in Figure 4.3.3 Design of MastersThe masters of the tooling are next developed by modifying the design of the actual tooling. It is known that the dimensional and surface accuracy produced by RP techniques can not meet the requirement for precise assembly in many cases. Therefore, an allowance of 4mm is applied to the surfaces that fit into the die holders for machining and polishing purposes.

Therefore, an allowance of 4mm is applied to the surfaces that fit into the die holders for machining and polishing purposes. Other modifications are optional like removing the step that is present in the actual tooling in order for easy releasing of the masters during the casting stage. Once this is proven, the 3D CAD models are converted into STL-formatted files ready to be input to the building stages. The die insert master isshown in Figure 5.Figure 5. The design of the master4. TOOLING PREPARATION4.1 Use of SLS527The steps involved in this approach (Flow A in Figure 2) are:4.1.1 Building of masters by SLSIn the SLS process [8], a thin layer of powdered material is deposited from feed cartridges over the working area by a roller. Within an inert nitrogen atmosphere, the powder is maintained at a temperature just below its melting point.

Within an inert nitrogen atmosphere, the powder is maintained at a temperature just below its melting point. A focussed CO2 laser beam is directed onto the powder, tracing out the pattern of the cross- section of the model to be built, and causing fusion of the powder. The process is thenrepeated with the next layer of powder until all layers are complete.The quasi-solid part is finally produced within a powder cake, which provides natural support for isolated parts. After slowly cooling down, the cake is gently removed. Finally, the parts are broken out and cleaned using a bead blaster machine. The required masters built on a Sinterstation 2500 (DTM Corp.) are shown along with the produced by other two routes in Figure 6.Figure 6. Masters made by SLS, CNC, and SLA4.1.2 Silicone rubber moldingOnce the masters are built, they are then used in a secondary process, or vacuum casting, to generate the non-ferrous forming tool.

Masters made by SLS, CNC, and SLA4.1.2 Silicone rubber moldingOnce the masters are built, they are then used in a secondary process, or vacuum casting, to generate the non-ferrous forming tool. Vacuum casting[9] uses silicone rubber that can be moulded around a master pattern to produce a cavity. Transparent silicon rubber (Product name: VTV 750 from MCP, Germany) is first mixed with the determined amount of catalyst (recommended ratio is 10:1) and de- gassed in a vacuum chamber. The de-gassed mixture is then poured around the master which is initially mounted to a sprue (typically with super-glue) and suspended within a casting frame. The casting frame with the master and uncured mixture is then further de-gassed at room temperature. This is done to avoid trapped air bubbles528which could cause mould surface defects if they happened to occur at the interface between the master and the silicon rubber.

This is done to avoid trapped air bubbles528which could cause mould surface defects if they happened to occur at the interface between the master and the silicon rubber. After the mixture has been properly de- gassed and poured into the casting frame, the whole frame with the master and mixture is then placed in a temperature-controlled oven and cured at 40(cid:176) C for about 5 hours.After curing, the mould is cut with a scalpel at the pre-determined parting line. Release of the master leaves a mould in two segments or more, which depends on the complexity of the master. It is easy to bend the segments together into a mouldwith tapes (as shown in Figure 7). The cavity constructed with the silicone rubber can be used to mould a variety of materials at relatively low temperature.

The cavity constructed with the silicone rubber can be used to mould a variety of materials at relatively low temperature. The biggest advantage of the silicone rubber molding is the material flexibility that facilitates taking out both the master and casting products from the mold no matter how complex the geometry may be.Figure 7. Silicon rubber mold assembled4.1.3 Aluminum epoxy castingThe casting material used in the study is the aluminum epoxy EP250 from MCP, which comes along with resin and hardener in two separate packages. This EP250 mixture is then poured into the silicon rubber cavity in a vacuum to avoid air bubbles. It is cured at a temperature of 50(cid:176) C for a duration of 6 hours after which the mould is separated to release the aluminium epoxy parts. The parts then go through a secondary curing at a temperature of 150(cid:176) C for 6 hours to strengthen its mechanical properties.

The parts then go through a secondary curing at a temperature of 150(cid:176) C for 6 hours to strengthen its mechanical properties. The finished parts are tested to have mechanical properties that are similar to those of aluminum. The EP250 parts have to be machined on a lathe machine to the nominated dimension and with the pin holes drilled before assembled into the die and punch holders. Three sets of the EP250 tooling by various approaches are presented in Figure 8.529Figure 8. The three sets of aluminum/epoxy tooling4.2 Use of SLABasically, the process is similar to that described in Section 4.1. The majordifference is that the master is built on an SLA250 machine, from 3D System, USA.The Stereolithography (SLA) process involves a control computer which reads in the build file in order to direct a focussed ultra-violet laser beam onto the surface layer of a vat of liquid photosensitive resin (photopolymer).

The laser energy causes the layer of resin to polymerise, so that it becomes solidified. Within the vat of liquid resin is a platform which holds the solidified resin layer. Once a layer is complete, the platform is lowered under computer control so that the liquid resin flows over the polymerised layer. Excess resin is removed by a blade, so that the depth of liquid resin remaining above the solidified layer is exactly equal to one layer thickness of the model being built. The polymerisation process is then repeated layer by layer. When the object has been completely built, the platform is raised above the vat of the resin to drain off the excess liquid resin that has adhered to the object.

When the object has been completely built, the platform is raised above the vat of the resin to drain off the excess liquid resin that has adhered to the object. The parts are post-cured in an UV oven before they are served as masters to create the silicone rubber molds, and eventually the EP250 tooling.4.3 Use of High Speed CNC MillingIt is also possible to work initially with the STL files and proceed onto high speed milling machining and vacuum casting as in flow C in Figure 2. The detail processes are almost the same as the approaches described previously. The masters of the forming tool in this case are fabricated using a high speed CNC milling machine. The experimental system is well developed so that it can operate on a spindle speed in excess of 70000 rpm to easily cut machinable plastic materials instead of steel. The feedrate adopted here is 400mm/s and the diameter of the ball-end tool is 1mm.

The feedrate adopted here is 400mm/s and the diameter of the ball-end tool is 1mm. The control programme for the system has a function to automatically generate tool paths based on STL-formatted files. Indeed, the STL files used in this project are uniform530for any of the approaches, which keeps the comparison independent of possible errors coming from data representations.4.4 Direct Use of RapidSteelThis approach can build the forming tools directly without going through secondary processes. RapidSteel 2.0, commercialized by the DTM Corp., is a stainless steel-based powder mixed with binder. By using the STL file of the actual tooling that has previously been developed, the Sinterstation 2500 System processes the powder. The result is called a 'green' part. The green part is then post-processed in a furnace where the binder is burnt out and the metal powders are bonded together through the traditional sintering mechanism.

The green part is then post-processed in a furnace where the binder is burnt out and the metal powders are bonded together through the traditional sintering mechanism. This cycle is named the sintering furnace cycle and it takes place at a temperature of 1120(cid:176) C. The part is then allowed to cool. At this stage, the product is referred to as a 'brown' part that exhibits durability, but is also porous in nature.After debinding, Bronze infiltration is the final step that will be conducted in the same furnace. This metal melts with elevated temperature and infiltrates the brown part via capillary action, resulting in a fully dense part. The part is then allowed to cool at a rate of 180(cid:176) C per hour. The fully dense component consists of 60% steel and 40% copper. The die inserts from RapidSteel are shown in Figure 9. And it is found that the RapidSteel part has a good machine-ability to be post-processed by milling, turning and polishing.

The die inserts from RapidSteel are shown in Figure 9. And it is found that the RapidSteel part has a good machine-ability to be post-processed by milling, turning and polishing. However, due to the existing of the copper and porosity in the final part, the RapidSteel product shows the weakness in strength, hardness, durability and other mechanical performance compared to the exact steel tooling.Figure 9. The RapidSteel die and punch insert (left) and the partially polished insert (right)4.5 Sheet Metal EmbossingThe tooling built from different approaches are machined and polished before531they are then assembled with the die holders and mounted onto the Turret Punch Press machine one after another (see picture in Figure 10). The embossing process is carried out on a 0.4mm thick aluminum sheet metal with the prepared tooling. Figure 11 shows the samples produced.Figure 10. The assembled toolingFigure 11.

The embossing process is carried out on a 0.4mm thick aluminum sheet metal with the prepared tooling. Figure 11 shows the samples produced.Figure 10. The assembled toolingFigure 11. Embossed sheet metal samplesResults show that the spring back effect for all embossed design on different tooling sets does not vary much and is of a small magnitude. Tearing of the aluminium sheet metal is experimented when the tools are fully closed or sits on each other. This may be due to the fact that the limiting depth is exceeded. It can be calculated that the theoretical maximum drawing depth allowed for a 0.4mm thick sheet metal is 1.2mm, or three times of the sheet thickness. As the actual depth reached in this case is measured to be 1.4mm, this exceeds that of the maximum allowed. Results like these are helpful to improve the design for the next iteration until no error occurs.In each case, no significant tool wear is visually observed after 10 stamping operation.

Results like these are helpful to improve the design for the next iteration until no error occurs.In each case, no significant tool wear is visually observed after 10 stamping operation. In order to keep the tools for the further study, the lifetime is not investigated by destroy testing.532It is observed that the high speed CNC milling/vacuum casting approach produces the best surface finishing of the embossed sheet metal. The worst of all the approaches is the SLS/vacuum casting approach. The result is obvious since the SLS tends to produce a master that is rougher than other processes. The quality of masters directly affects the final outcome of the tooling, and then that of sheet metal products.5. COMPARISON AND DISCUSSIONA total of three design iterations and several sets of prototypes are made in this case study. However, with more user experience and refinement to the tool, a higher success rate should be achievable.

However, with more user experience and refinement to the tool, a higher success rate should be achievable. In addition, a set of hard tooling is also subcontracted to Liau Ying Industries Engineering so that comparison between RT and conventional technique can be accomplished. By eliminating the unexpected iteration or interfering, a comparative study is being conducted on the designed tooling by different building techniques.The surface finish of each master and cast part is visually inspected and measured with a profilometer (Mitutoyo Surftest.301) before machining. The surface measured is the top face of the parts. Three measurements are performed on the target face of the masters and the cast parts, and the average of the measurements is then taken.Dimensions are also measured on the feature produced to determine the processes' accuracy.

As any parts of the features can be taken to be the reference dimension, the outer diameter is chosen as it is deemed to be the easiest to measure. The measured results are compared with the CAD design values and the differences yield the deviation. Both of the results on the surface roughness and dimensional deviation are summarised in Table 1.As one of the primary objectives of RT is to produce tools with the least cost and time spent, the time and cost comparison between various approaches are given in Tables 2 and 3 respectively. The leading time for each approach is summed up with all steps covered, assumed that the steps happen continuously without any delay. In other words, the machine occupying time is the major concern in the comparison. The cost calculation includes the material consumable only.Table 1.

However, the soft tooling, from either the non-ferrous or RapidSteel 2, manages to form functional parts using the relatively soft sheet metal like aluminium or copper material. Regarding the dimensional and surface accuracy, the result generated by CNC milling is the best because the stairstep problem is less serious in the milling process than in RP.There are few differences in time consumed between the four RT approaches. It takes only three or four days to produce the forming tool by whatever technique used in this case. Obviously from this project, the cost on producing the short-run tooling is extremely trivial compared to that on the steel tooling.Table 3.

Comparison of the different approaches by cost of tool developmentSLSHigh Speed CNC millingSLARapidSteel-2Conventiona l ToolingMaster (40g)Silicon Rubber (1.2kg) EP-250 Tooling (50g)$12$89$13Master (250x150x 50mm) Silicon Rubber (1.2kg) EP-250 Tooling (50g)$23$89$13Master (56g)Silicon Rubber (1.2kg) EP-250 Tooling (50g)$24$89$13Green Part (500g)Infil-trant (500g)Nitrogen/ Hydrogen$75$34$30Die insert $1500Punch $1500534SubTotal 20% Material Wastage$114 $22.8$125 $25$126 $25.2$139Total Cost$136.8*$150*$151.2*$139*Cost calculation does not include labor cost and machine depreciationThe following information is used for the cost calculation: • SLS Duraform • SLA Epoxy Resin SL5170 • CNC polymer (Cibatool BM5460) • Silicon Rubber VTV 750 • EP250 • RapidSteel2 • Brown Infiltrant= $2950/10kg = $4240/10kg = $450 per 1500x500x50mm = $1485/20kg = $2500/10kg = $2750/20kg = $1350/20kgIn addition, it is well known that RP techniques are able to generate freeform parts with almost no geometric constrain.

However, the high speed CNC milling approach may suffer the drawback when under cut is unavoidable or special fixtures are needed. The ability of complexity independency helps the research result widely applicable.Thanks to the time and cost compression technique, any amendment, either serious or trivial, is easy to make as the RSMP concept allows the modified design to be iterated within a few days until a satisfactory outcome is achieved.6. CONCLUSIONThe various approaches have been carried out to implement the concept of rapid sheet metal prototyping. The flexibility of CAD systems enables the optimization of the tooling needed for sheet metal manufacturing. With the further introduction of SLA, SLS, High speed CNC milling, and RapidSteel in the tool making stages, a short-run production of sheet metal product is possible with the lower cost and shorter leading time.

Other work will be done in the future to optimize the tooling design based on the calculation of data errors between the CAD model and reconstructed model generated by scanning the actual part. Finally, it will bring about more benefits from the close loop concept of the RSMP system.Acknowledgementsis a joint development in collaboration with Temasek Polytechnic and Nanyang Technological University's School of Mechanical and Production Engineering, with funding support from Singapore Totalisator Board.This projectReferences1. Jacobs, Paul F, “ Stereolithography and other RP&M technologies: from Rapid Prototyping to Rapid Tooling”, ASME Press 1996.535$30002. C.K. Chua and K. F. Leong, “Rapid Prototyping: Principles and Applications in Manufacturing”, John Wiley, 1997.3. Atkinson, David. “Rapid Prototyping and Tooling: A Practical Guide”, Welwyn Garden City, Herts: Strategy Publications, 1997.4. Daniel F. Walczyk and David E.

vancements in the SIS processBahram Asiabanpour1, Behrokh Khoshnevis2, Kurt Palmer2, Mehdi Mojdeh21Texas State University-San Marcos, TX 786662University of Southern California, Los Angeles, CA 90089-0193Reviewed, accepted August 13, 2003Abstract Selective Inhibition of Sintering (SIS) is a new layer-based rapid prototyping process. This paper reports the progress in research and development of the SIS process. Specific printer path generation method, experimentation with various powder and inhibitor materials, and systematic models leading to optimum performance given various factors affecting part strength, surface quality, and dimensional accuracy are presented.1- IntroductionSelective Inhibition of Sintering (SIS) is a new rapid prototyping process that like otherRP processes, builds parts in a layer by layer fabrication basis (Khoshnevis, B., et al, 2003).

The SIS developers hope that by successfully completing this process, they can present a very precise and inexpensive system to be used at the work tables of designers, modelers, and innovators to fabricate their product designs and ideas. The SIS process works by joining powder particles through sintering in the part’s body, and by sintering inhibition of some selected powder bed areas. The selected areas of the powder bed for sintering inhibition are wetted by a printer. The printer nozzle moves in the XY plane and prints the inhibitor liquid on the model slice pattern. The print pattern includes the part boundary and hatch pattern for each layer of the sliced CAD model.

The print pattern includes the part boundary and hatch pattern for each layer of the sliced CAD model. The Selective Inhibition of Sintering process includes the following major steps (see Figure 1):0- The inhibitor tank is filled with the inhibitor liquid and the powder feed tank is filled with polymer powder, the inhibitor tank pressure and printer voltage are set, and the printer home is found.1- The powder feed tank pushes the powder level upward in the amount of desired step(s). The build tank moves the powder level downward in the amount of desired step(s). The amount of lowering movement sets the layer thickness in the final part.2- The cylindrical roller from the source tank spreads the powder over the build tank. Theroller rotates clockwise and moves in the X direction.3- An X-Y printhead moves the print nozzle over the desired layer profiles.

Theroller rotates clockwise and moves in the X direction.3- An X-Y printhead moves the print nozzle over the desired layer profiles. The input file for the print pattern is the machine path file which is generated by the machine path generator system.4- The heat source attached to the moving part of the machine (including the printhead and roller) moves in the X direction and its temperature and movement create a sintered layer with an unsintered pattern in the selected area of the powder bed.Repeat from step 1 unless the part is finished.5- The model is removed from the sintered block and the post processing includingcleaning, and adding wax or adhesive to the fabricated part, completes the SIS process.25An Alpha machine was built for initial research studies. Following figure shows the SISmachine performance in detail:Halogen lampRecycle TankBuilder TankRollerFeeder TankBuilder Tank HeaterPrinterHeaterFigure 1.

Following figure shows the SISmachine performance in detail:Halogen lampRecycle TankBuilder TankRollerFeeder TankBuilder Tank HeaterPrinterHeaterFigure 1. Selective Inhibition of Sintering process steps262- MotivationBecause of different motivations, research on the SIS process was pursued in different directions. The first motivation for this research is due to the fact that without a machine path generation system, the SIS process will be limited to fabricating parts with very simple geometry. Also in many cases, lack of hatch path printing will make part separation from surrounding material impossible. The necessity for having such a system was the primary motivation for this research.

Also in many cases, lack of hatch path printing will make part separation from surrounding material impossible. The necessity for having such a system was the primary motivation for this research. Desirable properties for the new system include:An appropriate machine path (i.e., boundary path and hatch path) printing pattern that enables the SIS machine user to easily remove the fabricated part from the surrounding material.Machine path verification before sending the file into the SIS machine. • System capability to work in a wide variety of PCs with minimum computer memory. • Ability to process CAD models with any size and complexity. • Ability to fix or report the STL files errors. The second motivation is due to the fact that to reach an acceptable level of accuracy,strength, and surface quality in the fabricated parts, all effective factors in the SIS process should be identified and controlled.

A study for applying appropriate materials (powder and inhibitor) and identifying and optimizing effective factors on the quality (e.g., surface quality, part dimensional accuracy, and part strength) of the parts fabricated by the SIS process was necessary.3- Machine path generationThe newly developed machine path generator system (Asiabanpour, B. Khoshnevis, B.,2003) uses STL files with the ASCII format as input and works in two steps. The first step generates slices at each increment of Z by intersecting the XY plane with the facets within the part. The second step uses all individual unsorted intersection lines in each Z increment to form the contour. This step includes sorting the intersection lines, recognizing the closed loops and disconnections (STL file error), and generating the machine path (i.e.

This step includes sorting the intersection lines, recognizing the closed loops and disconnections (STL file error), and generating the machine path (i.e. boundary path and hatch path) and simulation files (Figure 2).Slice 1Subroutines: Point finding Point sorting Hatch printingSTL fileStep 1 SlicingSlice 2 : :Step 2 SortingMachine Path fileSlice nSimulation fileFigure 2.The two steps of slicing and machine path generationIn the first step the STL file is read as input. Slice files are then generated by executingthe slicing algorithm. Only the intersection of those facets that intersect current Z=z are27calculated and saved. In this step, one facet is read at a time from the STL file. Then the z , intersection lines of this facet with all XY planes for are calculated. The intersection lines are stored in the specified file for the associated z increment. This results in one intersection line on each XY plane.

The intersection lines are stored in the specified file for the associated z increment. This results in one intersection line on each XY plane. By repeating this process for all facets, a set of slices is generated. This algorithm only saves the data of one facet in the computer memory, therefore only a small amount of computer memory is needed, and there is no practical limitation on the model size. In this step each slice is saved in a separate file on the disk. This guarantees that step 2 is run much faster as compared with the case where all slices are saved in a single file. After the completion of the slicing process, a set of vectors becomes available in each z increment. These vectors are not connected and are not in sequence.££ zZZMin{ max,zMax {,zminABIn the machine path generation process, the software starts from one vector and tries to find the next connected vector to this vector.

Then it does the same for the newly found vector until it reaches the start point of the first vector (in the closed loop cases) or finds a vector with no leading attachment (in faulty STL files containing disconnections). To sort the vectors, the algorithm reads one vector at a time from a slice file and writes it to another file. This file is either path file, when vector is connected to previous vector, or temp file, when the vector is not connected to the previous vector. Therefore, the sorting process does not need a large amount of memory to sort the data, and there is no limitation on the number of vectors in a slice and on input file size. In addition, unlike many other slicing algorithms that cannot handle disconnections caused by faulty facets (Leong et al. 1996), this algorithm can generate a machine path even with disconnection errors in the STL file. At disconnection instances the system sends a message to a log file and turns the printer off and starts from a new vector.

At disconnection instances the system sends a message to a log file and turns the printer off and starts from a new vector. Also for each selected vector the possibility of hatch intersection points are investigated.At the end of the path generation process for one slice, the hatch intersection points are sorted and written into a machine path file. After the arrangement of all vectors in one slice (z increment), the process starts the arrangement of the vectors of the next slice. This process is continued until all vectors in all slices are sorted.Three subroutines are called in the machine path generation algorithm. The first subroutine is for finding the intersection points of every line and horizontal and vertical hatch lines. According to the algorithm, every line which is written to the output file is called in the following subroutine to find and save the intersection point(s) of the line with the hatch lines.

According to the algorithm, every line which is written to the output file is called in the following subroutine to find and save the intersection point(s) of the line with the hatch lines. At the end of formation of the boundary path, all intersection points are sorted in ascending and descending orders for every other line. The last step of the algorithm is setting the print/no print command for the hatch lines.Slicing and machine path generation algorithms have been implemented in the C programming language.

The last step of the algorithm is setting the print/no print command for the hatch lines.Slicing and machine path generation algorithms have been implemented in the C programming language. The software has been successfully tested for several medium and large STL files up to 200 MB on different PCs and laptops.Figures 3 and 4 show the algorithm implementation as presented by the path simulationmodule of the system.28C}}Figure 3.Visualization of the CAD model and the machine path in the AutoCAD environmentFigure 4.Visualization of two slices of the machine path in the AutoCAD environment4- Material selectionTwo types of materials are used in the SIS process: Powder and Inhibitor. Severalaffecting factors in the SIS process will remain constant by determining and using a specific powder and inhibitor for this process.

Severalaffecting factors in the SIS process will remain constant by determining and using a specific powder and inhibitor for this process. In addition to the technological requirements for the selected materials, these materials should also be non-toxic, easily accessible, and relatively inexpensive.4-1- Polymer selectionConducted experiments showed that many polymers (e.g., polystyrene, polycarbonate,polyester), depending on the desired properties of the final part, can be used in the SIS process. The only exception was polyamide (DTM), where the glass transition point is very close to the melting point. However, each polymer requires the appropriate setting. Based on conducted experiments polystyrene polymer was the most appropriate polymer for making high quality parts (e.g., sharp edges, good attachment between layers, and low deformation).

Based on conducted experiments polystyrene polymer was the most appropriate polymer for making high quality parts (e.g., sharp edges, good attachment between layers, and low deformation). Also, according to a discussion with a Degussa researcher, Degussa is able to make customized polymers (e.g., with specific melting point, shrinkage, deformation, and strength) based on customers’ needs. Therefore, many of the experimental limitations (e.g., part deformation) can be solved through material customization.294-2- Inhibitor liquid selectionAs the name implies, inhibitors are used to prevent the selected powder particles from sintering. Inhibition of sintering can be achieved through different phenomena. Regardless of the type of liquid, impact cutting and cooling phenomena are observed. Denser (heavier) liquids have more effective impact cutting. Liquids with high specific heat and evaporation energy constants have a more effective cooling phenomenon.

Denser (heavier) liquids have more effective impact cutting. Liquids with high specific heat and evaporation energy constants have a more effective cooling phenomenon. Several liquids including water, isopropyl alcohol, commercial cleaning agents, organic solvents, light lubricating oils, mold release agents, and various silicone liquids were tested as candidates for the sintering inhibitor. They demonstrated no obvious inhibition phenomena other than impact cutting and cooling.In addition, depending on the specific inhibitor chemical and physical properties, someother inhibition effects may be added to the list of inhibition phenomena. Inhibitor reaction with the polymer and separation phenomenon are two such effects.There are several materials that can dissolve polymers. Dichlorobenzene (C6H4Cl2) and butylbenzene (C6H5(CH2)3CH3) are two examples.

Dichlorobenzene (C6H4Cl2) and butylbenzene (C6H5(CH2)3CH3) are two examples. However, because of the toxic nature of the former and the destructive role of both in the printer head system, their application for the fabricated Alpha machine was impossible. In addition, polystyrene reacts with oxygen at high temperatures and in the presence of hydrogen peroxide, and hence inhibition occurs in the treated areas. However, it could not be demonstrated in the SIS process. Applying hydrogen peroxide as the inhibitor was successful for the low temperature sintering process (brittle or semi powdery parts), but for highly sintered parts the inhibition was not working at an acceptable level and removing parts from the surrounding materials was impossible.Another phenomenon for the inhibitor can be separation between polymer particles toprevent them from joining to each other. Applying this idea has had the most successful results.

Applying this idea has had the most successful results. In this phenomenon a soluble salt is printed as the inhibitor. After layer sintering and liquid evaporating, a thin layer of salt will remain in the selected area (Figure 5). At the end of the SIS process, this wall can be dissolved and washed away by water (Figure 6). A little alcohol is added to water to reduce the surface tension of the water. This enables water penetration into the solid block. Because of the weak powdery bounding between polymer particles in the selected area, the final part can be easily extracted from surrounding material. Potassium iodide (KI) has high solubility (135 gr/ 100 gr water) and a high melting point (635 o C). A saturated aqueous KI solution has been successfully used for many part fabrication experiments.30Salt SolutionDry Polymer PowderSolid SaltSintered Polymer PowderHeatingWet Polymer PowderPolymer-Salt MixtureFigure 5.

Water evaporation in the selected areaSolid SaltSintered Polymer PowderSintered Polymer PowderWashing in WaterPolymer-Salt MixturePowdery PolymerFigure 6. Washing the salt to extract the part5- Physical and mechanical part quality improvementSeveral properties (i.e., strength, surface quality, and dimensional accuracy) are important for the parts fabricated by the SIS process. Also, many different process operating conditions (factors) influence these properties. Statistical tools have been applied to improve the desirable properties of the parts fabricated by the SIS process. Initial attempts to simultaneously optimize all the responses for the SIS process produced several runs with missing data (immeasurable parts that were very brittle), and hence difficulties in applying statistical analysis tools.

Consequently, it was decided that a region of operating conditions that provide acceptable part strength should be identified prior to investigation of conditions that improve surface quality and dimensional accuracy. Part properties improvement started with an investigation of part strength using Response Surface Methodology (RSM) and a region of acceptable operating conditions for the part strength was found. Then, by using RSM, the impact of the factors on the final part surface quality and dimensional accuracy (for x,y, and z directions) is modeled. After developing a desirability function model, process operating conditions for maximum desirability are identified. Finally, the desirability model is validated.

After developing a desirability function model, process operating conditions for maximum desirability are identified. Finally, the desirability model is validated. Figure 7 illustrates the roadmap for this research.31Dimensional Accuracy Model Minimum deviation in X, Y, and Z directionsPart Strength Model Making Part with acceptable strengthMinimum shiftingOverall Optimization Desirability functionSurface Quality Model Making Part with optimum surface qualityFigure 7. Roadmap for the SIS process properties optimization5-1- Part geometryA sample block with dimensions 45¥ 20¥ 6 mm was selected as the base part geometryfor the experiments. Settings of the layer thickness factor produced specimen parts with differing numbers of layers; therefore, different machine path files were generated.Figure 8.

Settings of the layer thickness factor produced specimen parts with differing numbers of layers; therefore, different machine path files were generated.Figure 8. Base part5-2- Part strength testIn the previous research (Asiabanpour, B., Palmer, K., Khoshnevis, B., 2003) the transverse rupture strength test, the standard basis for measuring the strength of the low ductility powder metallurgy materials (ASTM B528), was used as the response measure for the experiments (Figure 9).32Force IndicatorPartFigure 9. Part breaking mechanismThen, a mathematical model of part strength versus effective factors was developed and region of acceptable operating conditions for the part strength (part strength > 50 oz) was found.5-3- Part accuracy responsesMinimum deviation between fabricated part dimension and CAD model dimension was selected as one of the part accuracy criteria. To measure the deviation, each axis (X, Y, and Z) was studied separately.

To measure the deviation, each axis (X, Y, and Z) was studied separately. For finding deviation of each axis, length (X), width (Y), and height (Z) values of the fabricated parts were measured using a caliper. Then, deviations from CAD model dimension were calculated as the error percentage. For example, deviation in the X axis wascalculated throughEX= X45 ¥45100, where X is the measured length of the fabricated part.By observation of the fabricated parts from the previous experiments it was noticed that some of the parts had been made with noticeable shift in the X axis direction. To minimize such a defective property, layer shifting in the X axis direction was selected as another part accuracy response (named shift_X). Shift_X (total shift in X axis direction between first and last layer) was measured by a caliper in mm (Figure 10).Shift XFigure 10.

Shift_X (total shift in X axis direction between first and last layer) was measured by a caliper in mm (Figure 10).Shift XFigure 10. Shift_X5-4- Surface quality responseLaser scanning systems are used to produce high precision surface quality evaluation.Mike Shellabear (1999) performed an extensive surface quality comparison between many commercial rapid prototyping systems. For surface quality measurement, he used a Zeiss Laser Scanning Microscope (LSM), which has a quoted accuracy of 0.2µm in X and Y directions and 0.6µm in Z direction.In this research, a rating system was used as a low cost alternative to evaluate the partsurface quality. For this purpose, ten sample parts with different surface qualities were selected33and assigned a rating between 1 and 10 to indicate the worst to the best surface quality.

For this purpose, ten sample parts with different surface qualities were selected33and assigned a rating between 1 and 10 to indicate the worst to the best surface quality. Then, all fabricated parts were rated by comparison to the selected sample parts.For precise rating, four boundary sides of each part are studied individually (Figure 11). Each side is rated based on three major properties: holes made of the accumulated salt (due to over printing), sharpness of the edges, and smoothness of the surfaces. At the end, an average value resulting from the 12 ratings (4 sides ¥ 3 properties for each side) is assigned as the part surface quality rating (Table 1).Surface holesEdge imperfectionFigure 11. Part breakdown for surface quality ratingTable 1. Part surface quality rating system PartSide 1Side 2Side 3Side 4Part gradeNo.

Part breakdown for surface quality ratingTable 1. Part surface quality rating system PartSide 1Side 2Side 3Side 4Part gradeNo. EdgeSmoothness Holes E S H E S H E S HSum/12K6854 5 65 5 592 65.5K+1 ……… .… .5-5- Factor identification & classificationBecause of the similarity between factors that were expected to be effective on surfacequality and factors that were expected to be effective on dimensional accuracy, a common list of eight factors was selected to be used in the new experiment. Based on previous experiences and preliminary experiments, eight factors are considered for statistically designed experiments. The list of factors includes two factors related to powder compaction, three factors related to inhibitor printing, and three factors related to sintering.

The list of factors includes two factors related to powder compaction, three factors related to inhibitor printing, and three factors related to sintering. Figure 12 illustrates the selected factors in the IDEF0 graph.34Machine feed rate Layer thicknessPrinter feed rate Printing frequency Printer pressureHeater feed rate Heater temperature Build tank temperatureCompacted powderPolymer powder1-Powder compactingPrinting time2- Inhibitor printingDroplet size, Frequency, & VelocityInhibitor Contour fileDroplets3- Inhibitor penetratingWet powder4 - SinteringFinal part Waste material5- InhibitingUnsintered materialBuild tank Feeder tank Roller ElectricityPrinter valve XY printer drivers Printer pressure pump ElectricityHeater ElectricityFigure 12.

IDEF0 hierarchy for the SIS process and selected factors for statistically designed experiments5-6- Design of experiments and analysis of resultExperimental design was implemented to characterize the SIS process in terms of howinput parameters affect the part surface quality and dimensional accuracy. The data set of designed experiments was used to develop mathematical models to show the responses behaviors versus the factors. Five mathematical models were developed for part accuracy (Ex, Ey, Ez, Shift_x) and surface finish measurement.5-7- Desirability functionReview of the developed regression models revealed the need for compromise in selection of factor settings. Each of the factors is involved in at least three of the models. Performance trends with respect to the factors are difficult to identify because of the inclusion of several two-factor interaction effects.

Performance trends with respect to the factors are difficult to identify because of the inclusion of several two-factor interaction effects. Obviously, it is plausible that an increase in the value of one factor might produce a more favorable result for one response but would produce a less favorable result for another response. It is necessary to balance the responses to achieve overall desirable performance. A multi-response optimization technique is needed. We use a desirability function for this purpose. The general form of the component desirability function for each response is shown in Figure 13. The response value labeled b in Figure 13 represents the target quality. This response value corresponds to the maximum desirability (d=1). In theory, the response values labeled a and c represent the limits of acceptability. For each response a desirability function was defined. The goal, of course, is to improve desirability for all of the responses.

For each response a desirability function was defined. The goal, of course, is to improve desirability for all of the responses. Then, a composite desirability function is generated by use of a geometric mean. To develop a mathematical model for the relationship between the factors and the composite desirability function, the individual response models were used to predict the response values over a grid of locations in the factor space. Finally, the composite desirabilty was calculated and a regression model is built for this new response.35Desirability10Responsea b c Figure 13. General form of the component desirability functionCanonical analysis and ridge analysis (Myers, R., and Montgomery, D., 2002) were usedto find the optimum point for the mathematical model of desirability function.Finally, the developed model was validated. To validate the model, 3 additional sample partswere fabricated.

To validate the model, 3 additional sample partswere fabricated. All five responses were measured for each part and average results were calculated for set of three parts. For each response, the individual response models were used to calculate 95% prediction intervals for the average of three values. All five of the actual averages fall within their respective prediction interval. These results indicate that the models are valid.6- Physical part fabricationSeveral 2.5 and 3D parts have been successfully fabricated by the SIS process using themachine and hatch path generation system. Figures 14 and 15 illustrate some of the fabricated parts.Figure 14.2.5D Parts fabricated by the SIS process36Figure 15.3D Parts fabricated by the SIS machine7- Vision for future researchResults from research indicate a very promising future for the SIS process.

SIS will be astrong competitive process against current RP commercial processes because of its ability to apply a variety of materials, as well as its accuracy, cost, speed, and ability to build functional part. However, many issues related to the SIS process should be studied and investigated. The multidisciplinary nature of the SIS process development opens a wide variety of areas for future work.Hardware issuesSeveral changes in the original SIS machine structure could result in processimprovement. Applying inkjet printer instead of a single nozzle printer makes the system much faster and more precise. Also, a larger build tank provides the ability to fabricate larger parts. In addition, improvement in heater system can cause more uniform layer thickness and consequently result in more accurate parts.CAD/CAM issuesCurrent machine path development is specialized for the single nozzle printer.

In case of applying inkjet printer, a CAD/CAM system is required to produce raster image, which includes boundary path and hatch pattern. Raster images can also provide data for multi-color printing to make multi-color parts. • Material issuesThis research has focused on a limited number of polymer powders. Based on required part properties, many other polymers should be usable. The SIS process is also expected to be applicable to metallic powders. We are currently investigating the use of iron and aluminum powders. These powders may require the use of inhibitors that capitalize on inhibition phenomena other than particle separation at the microscopic level. For example, inhibitors that produce chemical reactions with the powder, or those that produce particle separation at the macroscopic level.

For example, inhibitors that produce chemical reactions with the powder, or those that produce particle separation at the macroscopic level. Finding appropriate inhibitors to allow use of powders that produce a variety of part properties is a broad area of research.Powder compaction issuesThe fabrication of uniformly dense parts can be facilitated by compaction of the powderprior to sintering. This is especially true for metallic powders. However, the presence of the inhibitor in the powder layer complicates the description of compaction, specifically in terms of the dimensional stability of the layer. The extension of existing powder compaction models (ABAQUS 1997; Tran, et. al. 1993; Chenot, et. al. 1990) to the case of compaction for powder layers containing ‘interior walls’ of inhibitor material remains an open area for future work.37Powder waste reduction issuesVarious means should be devised and studied for powder waste minimization.

A combination of hardware and software approaches can significantly reduce powder waste in SIS and hence these approaches provide promising future research directions.Process optimization issueA systematic approach was explained in this thesis to optimize the process settings. However, all related factors were not considered in the statistically designed experiments. Consideration of other factors (e.g., material related factors) as well as application of other optimization methods (e.g., Taguchi) may lead to more realistic models.References: 1- ABAQUS/Standard User’s Manual, Version 5.7, vol. 1-3, Hibbitt, Karlsson, & Sorensen, Inc., Providence, RI, 1997.2- Asiabanpour, B., “An experimental study of factors affecting the Selective Inhibition of Sintering process”, Ph.D.

VANCES IN ELECTRON BEAM MELTING OFALUMINUM ALLOYST. Mahale*, D. Cormier*, O. Harrysson*, and K. Ervin**Edward P. Fitts Department of Industrial and Systems Engineering, North Carolina State University, Raleigh NC 27695** Department of Materials Science and Engineering, North Carolina StateUniversity, Raleigh NC 27695AbstractThe high thermal conductivity and melt pool optical reflectivity associated with aluminum alloys can pose significant challenges for direct-metal SFF processes. The use of SFF processes to produce aluminum parts is often not cost effective relative to CNC machining for simple geometries. However, the use of SFF techniques for aluminum alloys is justified for some applications such as aerospace forgings or high surface area heat exchangers. This paper describes recent progress in processing aluminum alloys using the Electron Beam Melting process.

This paper describes recent progress in processing aluminum alloys using the Electron Beam Melting process. Structure and properties will be discussed, as well challenges associated with high vapor pressure alloying elements such as zinc and magnesium.Introduction Interest in aluminum alloys for a wide variety of applications have increased significantly over the past two decades. Several aluminum alloys have been developed which have excellent strength to weight ratios. The relatively low melting points also make processing of these alloys more energy efficient. Aluminum alloys posses other thermal and electrical conductivity, ductility, desirable qualities such as high machinability and toughness.Research related to layered manufacturing of aluminum has largely been supported by the aerospace industry. The components used in this industry can be classified as low volume and high reliability.

The components used in this industry can be classified as low volume and high reliability. Annual demands for forged aluminum components in aging aircraft is commonly in the single digits. Forging dies for these components is often difficult to locate or is missing entirely. This has spurred considerable interest among the aerospace industry for SFF techniques capable of producing aerospace grade aluminum alloy components with forged or cast properties.Many of the commercial laser melt based systems suggest sintering of polymer coated metal powder to form a green part followed by a polymer binder burnout operation to form a denser part. After this the part can be further sintered to form a dense part. An alternative to this is infiltration of the sintered skeleton formed by after the polymer binder burnout and metal powder sintering. Sercombe et al. [1, 2] suggest improvements to the process chain for sintering and infiltration based manufacturing techniques.

Sercombe et al. [1, 2] suggest improvements to the process chain for sintering and infiltration based manufacturing techniques. A resin bonded Al-6061 part was first formed using a selective laser sinteringsetup. The infiltration was carried out by covering the part with alumina and magnesium powder and heating it to 540°C in the presence of nitrogen. The magnesium prevents the formation of an oxide inside the porous part and the nitrogen reacts with the aluminum to form AlN. The temperature of the furnace is raised to 570 C. The rigidity of the AlN skin permits it to be infiltrated by an aluminum alloy with a narrow melting range (Al-Si binary system is preferred due to its low viscosity when molten). At 570 C, the infiltrant melts and is absorbed into the porous part. The parts produced by this process can be considered to be comparable to investment cast parts of the same material. Souvignier et al.

The parts produced by this process can be considered to be comparable to investment cast parts of the same material. Souvignier et al. [3] in collaboration with Sercombe have experimented with developing a fused deposition modeling like process to make aluminum metal matrix composites. A PMMA based slurry loaded with aluminum alloy in combination with silicon carbide, carbon fiber and alumina powder is extruded through a syringe to form the base. A process similar to that described above was used to generate a dense part.Many researchers have also experimented with electron beam based melting for the production of dense aluminum parts. Matz [4], Taminger et al. [5] and Watson et al. [6] fed a raw material wire into the melt pool generated by the electron beam. Matz’s work does not make any specific reference to his group having worked on the development of the system for processing aluminum alloys.

Matz’s work does not make any specific reference to his group having worked on the development of the system for processing aluminum alloys. Taminger and Watson have described a low energy (15 kV) electron beam system aimed at in-orbit fabrication system. The deposition speed for this system varied between 400 mm/min to 800 mm/min while maintaining a constant wire feed rate. As expected, higher translation rate led to faster cooling and finer grain structure. Another experiment was carried out were the beam traverse speed was kept constant and the wire feed was varied between 3000 mm/min and 1500 mm/min. Higher wire feed rates led to faster cooling as the energy was distributed into melting a larger amount of material, thus producing small equiaxed grain structure.

Higher wire feed rates led to faster cooling as the energy was distributed into melting a larger amount of material, thus producing small equiaxed grain structure. The strength of parts built on this system seem to be comparable to 2219-O alloy but the surface finish of the part is dictated by the high weld pool width and height.Electron Beam Melting of Al 2024 2024 is an aluminum alloy whose primary alloying element is copper. It is known for its excellent fatigue performance, though it is not easily welded. 2024 Al was chosen as the first candidate for trials with the Electron Beam Melting (EBM) process due to the fact that the copper will not vaporize under vacuum substantially more or less than the aluminum. It was also readily available and relatively inexpensive.Feasibility tests for the processing of aluminum were carried out by placing a layer of aluminum powder inside a cavity approx. 0.75 mm deep in a steel plate.

0.75 mm deep in a steel plate. The experiments were cold started by directly scanning the cavity region at 10,000 mm/s between a current ramp from 0.5-3.5 mA that was covered through 100 repeats. The first experiment itself showed evidence of sintering with light balling occurring at the surface of the powder.Details of the initial feasibility tests are listed in Table 1. These tests were conducted to provide a starting point for the parameters prior to conducting experiments for producing a freeform part (particularly the powder preheating parameters).Table 1 Initial Feasibility Tests for 2024 AluminumPlateTemp.

in ˚CCold StartCold StartCold startCold start300˚CMelt cycle100 reps 10,000mm/s 0.5-3.5mA 0.5mm offset 100 reps 10,000mm/s 0.5-7mA 0.5mm offset 100 reps 10,000mm/s 0.5-4.0mA 0.5mm offset100mm/s 3.2mA 0.3mm offset100mm/s 3.2mA 0.3mm OffsetPre/Postheat NoneNonepod 100 reps 10,000mm/s 4.0-0.5mA 0.5mm offsetphd 100 reps 10,000mm/s 0.5-5mAphd 100 reps 10,000mm/s 0.5-5mA 0.5mm offsetObservationsSintering and minor BallingCracking followed by ballingofsinteredpartObjective: slowly decrease the heat as it would permit the material to flow after it has been melted to prevent balling. This did not help and balling was still observed Objective: Uniformly melt the balled Powder powder. segregated into two layers. Top layer balled and bottom layer sintered Objective: study the effect of a hot start.

segregated into two layers. Top layer balled and bottom layer sintered Objective: study the effect of a hot start. Observed that hot start did not have any significant effect on the quality of the melt300˚C100mm/s 1.8mA 0.3mm offsetPhd1 100 reps 10,000mm/s 0.5-5mA phd2 100 reps 10,000mm/s 2-5mAObjective: higher sintering of powder bed to provide a more to stable wettable prevent baling. It was observed that the balling occurred due to sputtering during the melt process.surfaceThe next set of experiments were meant to develop parameters for freeform fabrication of aluminum parts. A standard 100 mm square, steel start plate and a slice thickness of 0.07 mm was used for manufacturing aluminum 2024 samples.Approx. 100mmStart plateIIISamples (25mm sq x approx. 5mm tall)IIIIVPreheat regionFigure 1 Configuration for small sample tests of Al 2024As shown in Figure 1, four samples with different melt parameters were produced.

5mm tall)IIIIVPreheat regionFigure 1 Configuration for small sample tests of Al 2024As shown in Figure 1, four samples with different melt parameters were produced. The objective was to establish parameters that would produce parts that looked visually most acceptable. The details of each experiment along with qualitative observations are listed in Table 2.The next set of experiments was to make comparisons between the ultimate tensile strength of 2024 aluminum parts produced via the EBM system and values expected from parts produced by conventional manufacturing. Tensile test specimens as dictated by ASTM-E8 standards with suitable machining allowance were fabricated via the EBM system. Rectangular coupons were fabricated and then machined to final coupon dimensions. An important observation with respect these samples can be made as follows. The coupons frequently cracked during removal from the steel build plate.

An important observation with respect these samples can be made as follows. The coupons frequently cracked during removal from the steel build plate. The parts were removed from the start plate by tapping with a hammer and relying on the mismatched coefficient of thermal expansion between the aluminum and the steel substrate to separate the parts. This region obviously undergoes significant shear forces during cooling (mismatched CTE's) and part removal (tapping with a hammer).Table 3 shows the different processing conditions used to make four sets of coupons, and Figure 2 shows photographs of the resulting rectangular blanks prior to machining as well as close-up photos of the top and side surfaces.An approximate calculation of the average densities of the parts was made. The results are shown in Table 4. It is apparent that within the range of parameters used to produce these parts, the beam power had a greater influence on density than the beam velocity.

The results are shown in Table 4. It is apparent that within the range of parameters used to produce these parts, the beam power had a greater influence on density than the beam velocity. It is equally clear that considerable experimentation is still needed to get the density near 100%, although it bears repeating that this was a preliminary feasibility study intended to determine the relative ease with which the EBM process could produce geometric shapes to near net shape.Table 2 Freeform Fabrication Experimental Procedure for 2024 AluminumPlate Temp 480˚CPreheat 50 reps 10,000 mm/s 0.5-6mASample I 1.8mA, 100mm/sSample II 2.5mA, 100mm/sSample III 1.8mA, 250mm/sSample IV 2.5mA, 250mm/sObservation: Over melting observed in all samples 9mA, 500mm/s 550˚C50 reps 10,000mm/s 0.5-6mA, 0.5mm12.5mA, 500mm/s9mA, 800mm/s12.5mA, 800mm/sObservation: Kinetic energy developed by the melt pool pushes excess melt pool off the build region.

This material is very loosely sintered powder and is not damaging to the rake. The higher plate heating temperature also seems to be helping in holding the part more firmly to the plate. High scan rates (sample III & IV) prove to be more effective at reduce balling. Low scanning speeds (sample I & II) also seem to permit the beam to melt deeper into the part. Swelling of the surface was also observed in these samples Sample III had the cleanest finish; though overmelting seemed to be occurring in all the samples.550˚C30 reps 10,000mm/s 0.5-6mA, 1mm9mA, 1000mm/s12.5mA, 1000mm/s9mA, 1300mm/s12.5mA, 1300mm/sObservation: The objective of this experiment was to minimize the time spent preheating (also increases powder recovery) and improve the surface finish of the part by speeding up the melting and reducing overmelting. It was observed that the sintering during preheating can further be reduced.

It was observed that the sintering during preheating can further be reduced. The parts obtained during this experiment also showed strong evidence of overmelting.550˚C30 reps 10,000mm/s 0.5-5mA, 1.2mm9mA, 2000mm/s12.5mA, 2000mm/s9mA, 3000mm/s12.5mA, 3000mm/sObservation: Samples I & II exhibit swelling of the top surface and this is more prominent in sample II. Sample II & IV have a relatively flat top surface but they might have a small amount of porosity as small pinholes were observed in the top surface.550˚C30 reps 10,000mm/s 0.5-5mA, 1.2mm9mA, 2000mm/s12.5mA, 2000mm/s9mA, 3000mm/s12.5mA, 3000mm/sObservation: THESE EXPERIMENT WAS CONDUCTED WITH A SLICE THICKNESS OF 0.1MM IN ORDER TO SPEED UP THE PROCESS. It was observed that the surface of the samples was not as flat as those obtained in the previous experiment. The excess melt pool build up was also higher.

The excess melt pool build up was also higher. It was decided that for the current set of experiment it would be more appropriate to use a slice thickness of 0.07mm.Table 3 Tensile Test EBM Processing ConditionsStart Plate TempPreheatSample I-ISample I-IISample II-ISample II-II550˚C30 reps 10,000mm/s 0.5-6mA9mA, 3000mm/s12.5mA, 3000mm/s9mA, 2000mm/s12.5mA, 2000mm/sSAMPLE I-ISAMPLE I-IISample I-I sideSample I-I topSample I-II sideSample I-II topSAMPLE II-ISAMPLE II-IISample II-I sideSample II-I topSample II-II sideSample II-II topFigure 2 2024 Aluminum Tensile Test CouponsTable 4 Tensile Coupon Density ApproximationsDENSITY POROSITYSAMPLE I-I 2.273g/cc 18.25%SAMPLE I-II 2.512g/cc 9.635%SAMPLE II-I 2.425g/cc 12.787%SAMPLE II-II 2.519g/cc 9.383%The final step in the initial feasibility study was to produce a complex geometry shape with steep overhanging surfaces.

In order to do this, the complex turbocharger wheel shown in Figure 3 was fabricated, as was the set of lattice structures shown in Figure 4. The processing parameters used were same as those for Sample I-I in Table 3.Figure 3 Aluminum Turbocharger WheelsFigure 4 2024 Aluminum Lattice StructuresElectron Beam Melting of 7075 Aluminum 7075 aluminum is a very high strength alloy that is commonly used to produce structural aircraft components. On a percentage basis, the most significant alloying elements of 7075 aluminum are zinc, magnesium, and copper. These alloys are the basis upon which Mg(Zn,Cu,Al)2 precipitation strengthening takes place. It is well understood that melting metals under vacuum will result in some loss of alloying elements due to vaporization. At 10-3 mbar vacuum in the e-beam chamber, zinc vaporizes at roughly 300°C and magnesium vaporizes at roughly 375°C. In comparison, both aluminum and copper vaporize at roughly 1200°C.

At 10-3 mbar vacuum in the e-beam chamber, zinc vaporizes at roughly 300°C and magnesium vaporizes at roughly 375°C. In comparison, both aluminum and copper vaporize at roughly 1200°C. As the melting temperature of 7075 aluminum is approximately 600°C, it is to be expected that some zinc and magnesium will be lost due to their relatively high vapor pressures. This makes 7075 aluminum a much more challenging alloy to work with in a vacuum environment.In order to test the basic feasibility of processing 7075 aluminum via the EBM process, powder in the +325/-100 mesh size range was obtained from Valimet.

The same experimental layout shown in Figure 1 was used to test a very large range of processing conditions in which beam power, beam velocity, stepover distance, and many other parameters were varied.In order to get a feel for the magnitude of vaporization that is taking place, a representative sample of 7075 aluminum processed on the Arcam machine was sent to NSL Analytical in Cleveland, OH for chemistry analysis. Table 5 shows chemistry analysis for both the unprocessed 7075 Al powder purchased from Valimet as well as the 7075 Al samples processed in the e-beam machine. For reference purposes, the standard limits for 7075 Al are provided in the rightmost column. All elements of the e-beam processed samples fell within allowable limits with the exception (as expected) of zinc and magnesium.The good news from these results is that neither the zinc nor the magnesium is completely lost to vaporization.

This suggests that the rate at which melting and subsequent solidification takes place is rapid enough to preserve reasonable quantities of those alloying elements. Several potential solutions to this issue are currently being evaluated.Table 5 Vaporization of Alloying Elements In 7075 Aluminum Feasibility TestsElementAl Cr Cu Fe Mg Mn Ni Si Sn Ti V Zn ZrPowder Feedstock Result Balance 0.2% 1.59% 0.22% 2.22% 0.05% NR 0.12% NR 0.03% NR 5.64% NRArcam EBM Sample Result 93.9% 0.23% 1.75% 0.25% 1.53% 0.053% 0.016% 0.19% <0.001% 0.031% 0.011% 2.01% 0.022%Standard Requirement 87.1 – 91.4% 0.18 – 0.28% 1.2 – 2% Max 0.5% 2.1 – 2.9% Max 0.3% NA Max 0.4% NA Max 0.2% NA 5.1 – 6.1% NAIn order to qualitatively assess the degree of porosity in the EBM-processed samples, a number of samples were sectioned, mounted, polished, and examined via optical microscope. Both Figure 5 and Figure 6 show side section views of the samples.

Both Figure 5 and Figure 6 show side section views of the samples. The bottom surface shown in Figure 5 is the surface that was in contact with the steel plate upon which the samples were "grown". It is readily apparent that metallurgical bonding and mixing between the steel and aluminum took place. Figure 6 shows two different regions of a sample selected to illustrate the degree and nature of pores that are present in the samples. Each figure includes a 200 µm scale bar. For reference, the deposited layer thickness was 100 µm. The height of each image therefore represents approximately 10 layers of deposited thickness. It is readily apparent from the micrographs that there is complete metallurgical bonding between each deposited layer. Figure 6 also clearly shows grains spanning across multiple layers.Figures 6a and 6b show that the processes samples do not yet have full density.

Figure 6 also clearly shows grains spanning across multiple layers.Figures 6a and 6b show that the processes samples do not yet have full density. However, the achieved density at this stage of development has improved substantially from where it was at the start of this effort. An important factor to consider is whether the pores are due primarily to incomplete fusion of the powder particles or whether they are due to heat cracking. Of the two causes of porosity, incomplete fusion is by far the easier challenge to overcome. Figure 6b shows both spherical pores as well as pores that are more elongated and jagged in appearance. It is believed that the spherical pores are primarily the result of incomplete fusion. The long irregular pores are of greater concern, as they could be the result of heat cracking.

It is believed that the spherical pores are primarily the result of incomplete fusion. The long irregular pores are of greater concern, as they could be the result of heat cracking. The causes and remedies for both types of pores will be studied in the course of the full development effort.Figure 5 7075 Al Start Plate Interface (10X magnification)(a)(b)Figure 6 7075 Grain Structure and Voids (10X magnification)As of the present time, a small number of geometric shapes have been fabricated in the 7075 aluminum alloy to demonstrate that a basic shape including downward-facing surfaces can be produced. Figures 8 and 9 show simple mechanical shapes, while Figure 10 shows a more complex shape intended for use as a high surface area heat exchanger.

Figures 8 and 9 show simple mechanical shapes, while Figure 10 shows a more complex shape intended for use as a high surface area heat exchanger. Although considerable experimentation results are encouraging.remains,initialfeasibilityFigure 7 7075 Aluminum Sample PartFigure 8 7075 Aluminum Sample PartsTable 6 7075 Aluminum Lattice Heat ExchangerConclusions and Future Work This paper has described preliminary results of feasibility studies involving direct- metal freeform fabrication of 2024 and 7075 aluminum alloys using the Electron Beam Melting Process. In both cases, initial sets of processing conditions that yielded reasonably good geometric shapes including unsupported downward facing surfaces were located.

In both cases, initial sets of processing conditions that yielded reasonably good geometric shapes including unsupported downward facing surfaces were located. The primary issues to address in full-scale development efforts include elimination of porosity and control of vaporization of alloying elements.AcknowledgementsPartial support for these research efforts by NASA and DRS Technical Services isgratefully acknowledged.References1. Sercombe, T.B. and G.B. Schaffer, Rapid manufacturing of aluminum components. Science, 2003. 301(5637): p. 1225-1227.2. Sercombe, T.B. and G.B. Schaffer, On the role of tin in the infiltration of aluminium by aluminium for rapid prototyping applications. Scripta Materialia, 2004. 51(9): p. 905-908.3. Souveignier, C.W., et al., Freeform fabrication of aluminum metal-matrix composites. Journal of materials research, 2001. 16(9): p. 2613-2618.4.

vances in Functional Testing With SFF PartsAlan J. Dutson, Kristin L. Wood, Joseph J. Beaman, Richard H. CrawfordDepartment of Mechanical Engineering The University of Texas at AustinAbstractFunctional testing of SFF parts represents an exciting area of research in solid freeform fabrication. One approach to functional testing is to use similitude techniques to correlate the behavior of an SFF model and a product. Previous research at UT Austin has resulted in development of an empirical similitude technique to correlate the behavior of parts of dissimilar materials and geometry. Advances in the empirical similitude technique are presented in this paper. Sources of coupling between material and geometry characteristics that produce errors in the current empirical similitude technique are outlined. A modified approach that corrects for such errors is presented.

A modified approach that corrects for such errors is presented. Numerical examples are used to illustrate both the current and the advanced empirical similitude methods.I.IntroductionThe product development process can be abstracted into three broad categories: design; evaluation; and fabrication (see Figure 1). The design and evaluation activities are iterated until a satisfactory design (one that meets customer needs) is achieved. If the design and evaluation activities have been performed properly, then the product fabrication stage can be carried out without further changes to the design. If, however, the design has not been properly evaluated and refined, then product fabrication will yield unexpected results that can only be corrected through design changes.

If, however, the design has not been properly evaluated and refined, then product fabrication will yield unexpected results that can only be corrected through design changes. Since late design changes are extremely expensive and difficult to implement, it is essential that a design be thoroughly evaluated and refined before fabrication begins.Prototypes are a powerful and commonly used means of evaluating product designs. Prototypes can be either physical or virtual in nature. Virtual prototypes include computer models, simulations, virtual reality, etc. Physical prototypes can either be traditional prototypes, which are created through typical manufacturing processes and manual construction, or rapid prototypes, which are created through solid freeform fabrication techniques.Product DESIGNDesign EVALUATIONProduct FABRICATIONFigure 1.

Abstraction of Product Development Process.279Improvements in computer equipment and software tools have produced a trend in many industries of using more virtual prototypes and fewer physical prototypes. The primary reason for this trend is found in the relative benefits of virtual prototypes, which typically include low costs and the ability to quickly evaluate changes in product designs. Significant reductions in development time and costs have been attributed to the use of virtual prototyping techniques. However, virtual prototypes have not entirely replaced physical prototypes in most development processes.

However, virtual prototypes have not entirely replaced physical prototypes in most development processes. Reasons for the continued use of physical prototypes include the following:Physical prototypes are often better suited for certain types of product evaluations, such as ergonomics, proportions, customer feedback, etc.Physical prototypes capture physical phenomena that may have been overlooked in the virtual models.Testing of physical prototypes can be a valuable aid in the verification / refinement process for virtual models.The third item listed above indicates that physical and virtual prototypes can work as complements to each other in improving the design process. The goal of our research is to improve the way in which physical prototypes are used in verifying and refining virtual models.

The goal of our research is to improve the way in which physical prototypes are used in verifying and refining virtual models. The specific objective of this research is to reduce the time required to verify virtual models by performing functional testing on rapid prototypes instead of on traditional prototypes. Replacing traditional prototypes with rapid prototypes can result in significant reductions in product development cycle times.Functional testing has traditionally not been performed with rapid prototypes since material properties and part sizes that are available from RP technologies are rarely the same as those of the product. In other words, while the form of the rapid prototype may be the same as that of the product, the functional behavior is, in general, different. In order to overcome this limitation, similitude techniques can be used to correlate the behavior of the prototype with that of the product.

In order to overcome this limitation, similitude techniques can be used to correlate the behavior of the prototype with that of the product. The two similitude techniques that are considered here include the traditional similitude method (TSM), which is also known as dimensional analysis, and the empirical similitude method (ESM).II. BackgroundThe TSM is a similitude method that relies solely on dimensional information to correlate the behavior of two similar systems. The ESM uses empirical data to correlate the behavior of two systems. Both the TSM and the ESM are reviewed briefly in the following sections.2.1 Traditional Similarity Method, TSMThe field of dimensional analysis (herein referred to as the TSM) has developed over several centuries. The basic idea of the TSM is to create scale factors, based on the dimensions of system parameters, which can be used to correlate the behavior of two similar systems.

The basic idea of the TSM is to create scale factors, based on the dimensions of system parameters, which can be used to correlate the behavior of two similar systems. The first step in the TSM process is to recast the dimensional equation that describes the system into dimensionless form, as follows:ddg 1(,2,...,dn)⇒= 0fπππ ( N 2,,...,1)=0280(1)where dj, are dimensional parameters, πi are dimensionless products, and N < n. For two similar systems (a product, p, and a model, m) the dimensionless products can be represented asffππ ( , p p 1, ππ ( , m 1,m2,2,,...,,...,πNp ,πNm ,))==00or, in terms of a particular parameter of interest, say X, asπXp ,πXm ,==ffππ , ( p p 1, ππ , ( m 1,m2,2,,...,,...,πNp ,πNm ,− 1)− 1)π = Xp ,πππ =for all i = For these two corresponding systems, the TSM states that 1, 2, (cid:133), N-1.

(Many references exist which show systematic derivations of dimensionless parameters from sets of dimensional parameters. See for example Barr, 1979 or Langhaar, 1951.)ifXm ,ip ,im ,ππ =for all i = 1, 2, (cid:133), N-1) are said to be well-scaled, while systems which do not satisfy the constraints are said to be distorted. Many sources of system distortion exist which can produce errors in the TSM approach. If, for example, one of the dimensionless products contains a parameter that is constant in the model system but variable in the product system, then the system becomes distorted and the TSM yields inaccurate results.Two systems which satisfy the TSM constraints (ip ,im ,2.2. Empirical Similarity Method, ESMThe empirical similarity method provides a means of correlating distorted systems. The fundamental concept of the ESM is shown in Figure 2.

Empirical Similarity Method, ESMThe empirical similarity method provides a means of correlating distorted systems. The fundamental concept of the ESM is shown in Figure 2. Unlike the traditional method, which relies solely on dimensional information for correlating systems, the ESM is able to correlate distorted systems by utilizing empirical data from a simplified specimen pair. The model specimen (ms) is a geometrically simplified version of the model, while the product specimen (ps) is a geometrically simplified version of the product. The ESM uses measured values from the model specimen, the product specimen, and the model to predict the behavior of the product. The basic assumptions of the ESM are as follows:1. The model and the model specimen can be tested to determine the state variation caused by pure geometric changes (G).xmsMxps = M xmsModel SpecimenGG’Product Specimenxm = G xmsM’xp = M’ xm = G’ xpsModelProductFigure 2.

Empirical Similarity Method. Adapted from (Cho, 1999).281(2)(3)2. The model specimen and the product specimen can be tested to determine the statevariation caused by pure non-geometric (material and loads) changes (M). The state of the product can be predicted by multiplying the state of the model by M’ or by multiplying the state of the product specimen by G’, as shown in Figure 2. A basic assumption of the ESM is that M = M’ and G = G’ (Wood, 2002), or that M and G are independent.the shows applicability of ESM with respect to the TSM and direct product testing. We claim the ESM is a more accurate approach, in general, than the TSM. We also claim the ESM is a better approach for correlating systems with complex geometry whose governing parameters may not be well known, as required by the TSM. However, the range of application of the ESM, which is represented by the boundary lines in Figure 3, has not yet been clearly established.

However, the range of application of the ESM, which is represented by the boundary lines in Figure 3, has not yet been clearly established. An evaluation of the current ESM boundaries, as well as a means to extend those boundaries, follows.Figure3qualitativelyy c a r u c c A d e r i u q e RESMDirect Product TestsTSMESMComplex GeometryFigure 3. ESM vs. TSMIII. Evaluation of the Empirical SimilitudeMethodThe ESM theory is valid as long as M = M’ and G = G’, as shown in Figure 2. Conditions that can cause M ≠ M’ and G ≠ G’, which are termed specimen distortions, are summarized in Figure 4. If the ESM can be set up with consistent scaling and with consistent material propertiesS ources of Error in ES M ApproachInconsistent S calingModel / product scaling vs. model specimen / product specimen scalingModel specimen / model scaling vs.

model specimen / product specimen scalingModel specimen / model scaling vs. product specimen / product scalingInconsistent Material PropertiesWithin parts (molding, RP)Between parts (concrete, RP)Distinct Geometric ConfigurationLarge vs. small deflectionInconsistent Material Behavior for Different GeometriesDistinct Operating Point along a Nonlinear M aterial Property CurveNonlinear σ-ε curve; E = f(load)T emp-dependent thermalconductivity; k = f(T )Viscoelastic material; δ = f(time)Functionally Coupled ParametersDamping coefficient (functionof material and geometry)Figure 4. Sources of Error in the ESM.282in the model and product families, then the only source of specimen distortion will come from inconsistent material behavior in the model or product family.

Sources of Error in the ESM.282in the model and product families, then the only source of specimen distortion will come from inconsistent material behavior in the model or product family. It is important to notice the subtle distinction that is made between material properties and material behavior: material properties refer to the global properties of the material; material behavior refers to the response of the material under some specific loading condition. For example a stress-softening material will behave differently – either “flexibly” or “stiffly” – depending on its specific operating point along the nonlinear stress-strain curve.The ESM error that results from inconsistent material behavior is perhaps more subtle and difficult to anticipate than the other classes of specimen distortion. Three finite element studies are used to illustrate this type of specimen distortion.

Three finite element studies are used to illustrate this type of specimen distortion. Each study illustrates one of the three subclasses of material behavior distortion that are listed in Figure 4. All three studies involve the deflection of a cantilever beam under an applied load at the tip. Each study involves some type of distortion between the model and the product (otherwise the TSM would be used rather than the ESM).Product SpecimenModel Specim enbe product evaluated is a cantilever beam with five holes along the length. The ESM setup is shown in Figure 5. In each study, three different hole diameters are considered for the product beam (small 0.15” holes with diameter, medium holes with 0.25” diameter, and large holes with 0.35” diameter). Each beam is modeled with linear shell elements (S4R elements) using ABAQUS software. Large deflection effects are considered in each study.

Each beam is modeled with linear shell elements (S4R elements) using ABAQUS software. Large deflection effects are considered in each study. Beam deflection is monitored and recorded at ten equally spaced load increments.Theto6.884.000.860.125" thick0.215" thickModelProduct6.884.000.860.667" 5XDiamproduct 5XDiammodel 5X 0.215" thick1.147" 5X0.125" thickFigure 5. ESM Setup for Finite Element Studies.3.1. ESM Study 1: Linear vs. Nonlinear Material PropertiesThe first ESM study involves a linear stress-strain curve for the product family and a nonlinear stress-strain curve for the model family. The constant value of Young’s modulus for the product family is 10,150 ksi (equal to that of aluminum).

The constant value of Young’s modulus for the product family is 10,150 ksi (equal to that of aluminum). The variable value of Young’s modulus for the model family is defined with a Ramberg-Osgood curve, which is described by the following equation (see ABAQUS, 2001):ε=σE   1+ α  σ σ 0  n−1   where σ = stress, ε = strain, E = Young’s modulus (defined as the slope of the stress-strain curve at zero stress), α = “yield” offset, σ0 = yield stress, and n = hardening exponent for the “plastic” (nonlinear) term. By using different parameter values in equation 4, three different sets of2830.50"0.50"(4)increasing properties with material degrees of nonlinearity were defined (see Table 1). Evaluating three geometric cases for each of the three different material properties gives a total of nine cases for this study.

Evaluating three geometric cases for each of the three different material properties gives a total of nine cases for this study. the ESM In each case, prediction of beam deflection is compared to the actual beam deflection, and a percent error in the ESM prediction is calculated. A plot of the results is shown in Figure 6 (note that the degree of model distortion corresponds to the material number). The error at “0” model distortion, which corresponds to a well-scaled system, is assumed to be zero.Table 1. Parameters for Ramberg-Osgood Curve. n 2.0 3.0 2.5αααα 0.40 0.43 2.00σσσσ0 (ksi) 3 3 3E (ksi) 150 320 500Material 1 Material 2 Material 3Because the model material has a nonlinear stress-strain curve, a in geometry produces a change different effect in the model family than it does in the product family. This source of ESM error can be understood more clearly by plotting the maximum stress for the various geometric configurations.

This source of ESM error can be understood more clearly by plotting the maximum stress for the various geometric configurations. Figure 7 contains such a plot for material 2. Notice that as the holes get bigger, the maximum stress (which occurs at the stress concentration around the first hole) increases. As the maximum stress increases, the model material behaves in a more flexible manner (i.e. the effective value of Young’s modulus decreases) while the product material remains the same (with a constant value of Young’s modulus). Since the material behavior is dependent on geometry, the ESM assumption is violated and errors result.Errors in ESM Prediction30)%(Degree of Change in Geometric ShapeGeometry 1 (Small Holes)25r o r r E r i c ,20Geometry 2 (Medium Holes)15M S E e g a r e v A10Geometry 3 (Large Holes)500123Degree of Model DistortionFigure 6. Results for ESM Study 1.3.3.

Maximum Stress for Various Geometries.284Table 2. Beam Lengths.Length 1 Length 2 Length 3 Length 4 Length 5Length (in) 4.00 6.88 8.00 12.0 16.0ESM results. The first length value represents a model beam that is “too short,” and the last three length values represent model beams that are “too long.” The results of the study are shown in Figure 8.)%(r o r r E r ic,M S E e g a r e v A35302520151050Errors in ESM PredictionChange in Geometric Shape10123Degree of Model DistortionGeometry 1 (Small Holes)Geometry 2 (Medium Holes)Geometry 3 (Large Holes)Figure 8. Results for ESM Study 2.The ESM error is caused by thefact that the length distortion in the model family produces a different geometric configuration in the model family than in the product family.

Because of large deflection effects, a change in geometry in the product yields a different change in beam deflection than a corresponding change in the geometry of the (already highly deflected) model.3.3. ESM Study 3: Isotropic vs.Orthotropic Material PropertiesErrors in ESM PredictionThe final ESM study involves isotropic material properties in the orthotropic product material properties the model family. Both the tensile modulus and the shear modulus in the z-direction (out of the page in Figure 5) of the beam were distorted. Since deflection due to shear is dependent on both the shear modulus the beam geometry, the ESM assumption is violated, and errors result. Results of this study are shown Figure 9.familyand inand)%(r o r r E r i c,M S E e g a r e v A141210864202Change in Geometric ShapeGeometry 1 (Small Holes)Geometry 2 (Medium Holes)Geometry 3 (Large Holes)1012Degree of Model DistortionFigure 9.

Figure 11 shows the improvement in ESM prediction accuracy for the “large holes” case of Study 1 using both 1st order and 2nd order polynomials. Similar improvements in prediction accuracy were also realized in studies 2 and 3.the higherthe degree ofModified ESM Predictions)%30(r o r r E M S E e g a r e v A252015105ESM1st Order Correction2nd Order Correction00123Degree of Model DistortionFigure 11. Modified ESM Results for Study 1.V. ConclusionsThe ESM has shown significant improvements over the TSM in predicting product performance when distortions between the product and the model exist. While the ESM typically produces highly accurate results, several sources of specimen distortion, which cause the ESM assumptions to become invalid, have been identified. The studies presented in this paper illustrate the ESM prediction error that can result from various types of specimen distortion.

The studies presented in this paper illustrate the ESM prediction error that can result from various types of specimen distortion. A modified ESM approach, which uses intermediate specimen pairs to capture the change in the transformation matrix as the geometry changes, has been shown to improve ESM predictions.Further areas of research include expanding the results presented here into other areas, such as heat transfer applications. In addition, further research is needed to be able to quantify the change in the transformation matrix when the geometry change involves different geometric shapes. Finally, an overall approach for functional testing with rapid prototypes is needed which indicates which method (TSM, ESM, modified ESM, or direct product testing) is appropriate for different situations.ReferencesABAQUS/Standard User’s Manual, version 6.2. 2001. Hibbitt, Karlsson & Sorensen, Inc. Barr, D. I. H. 1979.

vances in Selective Area Laser Deposition of Silicon CarbideJ. JI: Tompkins, R.. Laabi, B. R. Birmingham, H. L. Marcus Centerfor Materials Science and Engineering The University of Texas at AustinAbstractSelective Area Laser Deposition (SALD) is a Solid Freeform Fabrication (SFF) technique which uses a scanning laser beam to produce solid material by locally decomposing a gas precursor. In this work, a focused C02 laser beam strikes a substrate in the presence of tetramethylsilane (TMS) or diethylsilane (DES), producing silicon carbide objects with high density and no binder phase. Recent investigation has yielded growth rates up to 2.7mlnJmin in the beam area, and has eliminated previously noted contamination of the optics by a byproduct which mass spectroscopy identifies as silicon dioxide. This paper reviews a cause of non-uniform growth and delTIOnstrates the addition of hydrogen and reduced scan speeds to lTIake lTIultilayer parts.

This paper reviews a cause of non-uniform growth and delTIOnstrates the addition of hydrogen and reduced scan speeds to lTIake lTIultilayer parts. In addition, it presents a lTIethod for in-situ measurement of height of deposited material.IntroductionSelective Area Laser Deposition [1] describes a rapid prototyping process in which amoving laser beam causes localized decomposition of a precursor gas, creating solid material. A computer containing specifications for a three dimensional object controls the lTIotion of an x-y table, which moves the beam relative to a substrate surface. The moving area heated by the beam dictates where lTIaterial will be added, building up solid material (metal, ceramic, glass, etc.) in the desired shape for either prototype or structural parts.This paper reviews recent progress in SALD employing tetramethylsilane (Si(CH3)4) and diethylsilane (H2Si(C2H5)2) to produce silicon carbide.

Pyrolysis oftetramethylsilane has been investigated by Figueras et al in the growth of silicon carbide films [2]. Tetramethylsilane deCOlTIpOSeS according to the following overall reaction:The reaction proceeds exothermically, with an enthalpy change of -73.75 kcal/mole [3].We also considered an alternative precursor, diethylsilane. The molecular formula fordiethylsilane suggests the following overall reaction:Diethylsilane offered the possibility of lower deposition temperatures, as both the hydrogen and ethyl groups display less electronegativity and should ease decomposition relative to the methyl groups in tetramethylsilane.412A previous paper reports the fabrication of single layer silicon carbide objects fromtetramethylsilane by Selective Area Laser Deposition [4]. Thennal shock and uneven growth posed the greatest obstacles to the production of larger, more general shapes.

Thennal shock and uneven growth posed the greatest obstacles to the production of larger, more general shapes. The investigation of rod growth with a stationary beam produced growth rates of up to 1.6 mm/min. The following describes the most recent advances in producing more complex, multilayer silicon carbide objects, including increased growth rates, more even growth, and reduction in the effects of residual oxygen on the process.ApparatusThe equipment setup (Fig. 1) varies somewhat from early SALD work [5]. Vapor drawnfrom a flask containing liquid tetramethylsilane (vapor pressure 755 Torr) or diethylsilane (85 Torr) at ambient temperature fills a vacuum chamber. A computer controlled x-y table supports a CO2 laser and associated optics which produce a focused beam passing through a ZnSe window into the chamber. A firmly packed powder bed within the vacuum chamber serves as a substrate.

A firmly packed powder bed within the vacuum chamber serves as a substrate. Movement of the table scans the beam over selected areas of the substrate, whereLensLaseryxManometerWindowPrecursor Gas InletVacuum PumpH2 InletLiquid PrecursorFig. 1 SALD System Schematicpyrolysis of the tetramethylsilane or diethylsilane produces silicon carbide. In a departure from the earlier system, we have added a small fan with a brushless DC motor inside the chamber to provide forced convection to the deposition area similar to that in a flowing gas system.Experimental ProcedureThe vacuum chamber underwent multiple cycle purges with nitrogen to base pressures of less than 10-2 Torr. The system did not undergo a bakeout between trials. The substrate consisted of finnly packed 16um silicon carbide powder.

The system did not undergo a bakeout between trials. The substrate consisted of finnly packed 16um silicon carbide powder. Powder provides greater absorption of the beam, as well as less conduction of heat away from the deposition point, assisting in initiation of the SALD process.413Precursor gas pressure ranged from 20-140 Torr for tetramethylsilane. The upper limitreflects previous problems experienced with fogging of the ZnSe window by decomposition products at higher pressures. An equilibrium vapor pressure of 85 Torr at room temperature limited diethylsilane pressures to 85 Torr or less. Where indicated below, we also added chromatograph grade hydrogen gas at pressures up to 100 Torr.Two successive ZnSe lenses acted to focus the beam from the CO2 laser, producing an estimated 400um spot size (Ve2).

Previous measurements of absorption by tetramethylsilane provided a pressure/absorption correction factor to achieve a constant intensity at the substrate surface of six watts regardless of pressure. Trials for diethylsilane represent the first investigation for this precursor, and considerations outlined below prevented the acquisition and use of correction curves for uniform power at the surface.Scanning ofthe laser beam typically took place at 80-100 um/sec. We have notthoroughly investigated a full range of scan speeds, and future work will address the determination of an optimal scan speed. All scanning patterns consisted of parallel lines with a spacing of 100 urn.ResultsAs seen in Fig. 2, we observed growth rates as high as 2.7 mm/min. The previousmaximum rate of 1.6mm/min occurred at 80 Torr ofTMS. Rates had previously decreased at pressures over 80 Torr, due in part to clouding of the ZnSe window by precipitation of a reaction byproduct.

The difficulty of exactly reproducing the positioning of the deposition point relative to the small fan used in these experiments414probably accounts for much of the scatter observed resulting in fluctuations additional source of variation.output power asMode hopping in the laser,represent anAirflow from the fan offers an additional benefit:convection from byproducts upward along as the absorption of the beam. leaving the deposition site, greatly reducing clouding of theheated area carried a relatively undiluted stream of heatedbeam path, ultimately striking surfaces on optical components suchwindow.these surfaces, a solid materialfromto stream waste gasgas,addition of forced convection dispersedpath.Trials with diethylsilane did not show Observation of the substrate showed no heating of the settings (30 Watts).

Only at pressures below 10didatsurface, and at 10 torr a small amount of growth of very strong absorption of the 10.6 urn wavelength, the laser could not deliver the desired six watts to the ""1"",....."",."'" at pressures below 10 Torr suggest that solving the to serve as a suitableissue of growth rate represents a critical concerntechnology. The data shown above for tetramethylsilane growth rate over that previously reported. Fig.3 shows an a.V<~1'V\1",,1a. of a object. current state. The three horizontal appendages contain a carbide. This tnaterial has a durable and solid feel, but require further investigation.nearly a"""'..... ,......,........,..,scale of the object, approxiInately onethe capabilitythick layer of deposited silicon density and strength datamentioned above, previous work had revealed a chamber due to a byproduct of the depositionwith contamination Mass """"'..--+....£'1.<:',,,£'1.1"...

Toptics in electron impact ionization identified the solid byproduct precipitating on the reactor walls as well as on the ZnSe window as pure Si02. Based on eliminated the contamination problem: drying the substrate deposition, redesigning the precursor flask to eliminate grease, and adding hydrogen which adsorbs onto dramatically reduced the formation hours of deposition have not resulted in any visible However, hydrogen slows the deposition rate significantly: by approximately a factor of five.

Future work will investigate hydrogen needed as a function oftetramethylsilane pressure to prevent silicon dioxide TA...·1'V\<:.+.r.....information, threehave hours prior to to oxygen-bearing vacuumvacuum forHydrogen acts to getter any availableto formchamber walL Adding hydrogen tosilicon dioxide,involving up to optical components.Torrreduces the ratemtntmmnofAlthough silicon carbide generallyexcellentshock specificallyfor a ceramic ofmaterial, thermal shock also presented difficulties in multiple layers of materiaL Scan rates of 400 urn/sec produced solid single suitable in a furnace. the 100 um/sec reduces this problem.

the 100 um/sec reduces this problem. Two separate contributions initial layer of material has greater thickness and thereforetheanalysis by XRDS, SEM, scanning auger spectroscopy, or for annealing experiments shock caused by Reducing the scan speed to 80-deposition of a second layerseveremoving thermal gradient, fracturing the originalaccount for this: strength~second,4153created in 3 hours 15 minutes from tetralnethylsilane.slower scan rate induces a proportionately lower thennal shock on the material. Thennal shock under current conditions no longerwith growth of multiple layers of silicon carbide.The requirement to heatdeposition surfaceabove with laser radiation asopposed to heating the substrate from below as in conventional in SALD: any projection above underlying solid, causing it to retain more of the incident energy. This causes it to grow faster, and an unstable situation exists which we refer to as preferential growth.

This causes it to grow faster, and an unstable situation exists which we refer to as preferential growth. An instability in deposition rate identified during rod growth cOlnpounds the problem by actually initiating projections along the deposition beam, an unexpected periodic variation appeared in rod diameter, referred to hereafter as "beading". Fig. 4 shows an extrelne case of this. This instability Inay also exhibit itself in the growth of single or Inultiple layers of material, but became directly observable during the growth physical phenomena. Birminghaln [6] identified small of rods. Beading results from periodic fluctuations in the C02 and later showed that the associated shift wavelength caused by Inode hopping created a dratnatic shift in absorption by the gas.

Beading results from periodic fluctuations in the C02 and later showed that the associated shift wavelength caused by Inode hopping created a dratnatic shift in absorption by the gas. Further, he showed that reflectivity of the silicon carbide powder which very closely withcreates a unique challenge planar surface of the material suffers less conduction into theDuring growth rate measurements using a stationarydeposit also varied sharply in the regionCO2 laser's mode hopping occurred. The period of the beading events correspondsestablished period of mode hopping in thelaser.We elnploy two approaches to minimize the problem of localized growth: reducing thescan reducing heat flow intoreduces the thermal gradient between the substrate surface and the underlying Inaterial,substrate. Also, the addition of hydrogen increases the rate of416Fig.

Also, the addition of hydrogen increases the rate of416Fig. 4 Example of "beading" occurring during rod growthetching (reverse reaction) of silicon carbide at the higher temperatures experienced by projections on the surface. Although these approaches reduce preferential growth, the answer ultimately lies in the use of precursors with lower deposition temperatures to further reduce the thermal gradient in the material, and hence conduction into the substrate. Such precursors exist but require cOlnplex and critical handling procedures. In addition, SALD requires a laser with a wavelength appropriate to the chemical and physical properties of the material produced.In-situ deposition thickness measurement approachDevelopment of an in-situ deposition thickness measurement technique is critical tocontrol the height of the different shapes built up by the SALD process.

One lnethod consists of using the scattered energy of the He-Ne laser focused beam to make an image of the spot. This method is easier than the interferOlnetric lnethods and does not require polished surfaces, nevertheless its sensitivity may approach that of lnedium accuracy interferometry. This system is shown schematically in Fig. 5.Measuring principlesThe He-Ne laser beam is focused onto the surface of the growing object by the lens Ll. When the surface is exactly at the focal plane of the lens, a large fraction of specular reflection light is backscattered and collected by the lens L2 to project the spot image onto the position sensitive detector (PSD). When the deposit grows, the spot image moves along the detector which is an optoelectronic sensor that provides continuous position data of a light spot traveling over its surface. An aperture is used in front of L2 to improve the resolution by decreasing the417spotsize.

An aperture is used in front of L2 to improve the resolution by decreasing the417spotsize. Knowing the displacement d along the detector we can estimate the growth G by the following formula:IG=(XsinaIX' )*d( XsinalX') is a ratio associated with the position ofthe lens L2.C02 LASERAPERTUREPSDL2He-Ne LASER!~XElectronic circuitCHAl\fBERGrowthX,Y SCANNING TABLEComputer Growth=f(X,Y)Figure 5 Schematic of the deposition thickness measurement systemThe DetectionWhen a light spot falls on the PSD, an electric charge proportional to the light energy is generated at the incident position (Fig.6). This electric charge is driven through the resistive p layer and collected by the electrodes.

This electric charge is driven through the resistive p layer and collected by the electrodes. Since the resistivity of the P-Iayer is uniform, the photocurrent collected by an electrode is inversely proportional to the distance between the incident position and the electrode.418The following formulas give us an estimation of the incident position depending on thetwo outputs II and 12.II = IO(L-d)/2L 12 = IO(L+d)/2LINCIDENT LIGHTOUTPUT 11~IOUTPUT 12PhotocurrentP-LA YERN-LA YERi r----L --O+O---L -~10==11+122L = D istall ce betweell th e tw 0current leadsFigure 6 Schematic of the position-sensitive detectorOne of the most important problems that we met is the specular reflection on the rough surface of the silicon carbide, which causes a variable incident intensity of the spot image ( IO variable ).

To avoid that problem an operating electronic circuit was built in order to calculate the ratio: (12-Il)/(I2+Il) = dIL which is independent of the reflected light intensity.1111-12SPOT IMAGEP S Df-----JOPERATING CIRCUITANALOG DIVIDER",(11-12)/(11+12)---tI1211+12Finally we have:G = (Xsina/X')*L*(I2-II )/(Il +12)I419ResultsInitially, different measurements were made by changing the ratio (XsinalX'). The spotwas directly projected onto a fluorescent screen. The smallest displacement d resolvable was d = 0.5 mm which corresponds to 30 f..tm of growth with a ratio of (XsinalX') = 0.06. The PSD will be used to improve the position resolution of the spot image. A detected spot displacement of 150 f..tm is possible with such a detector which corresponds to a resolvable 10 f..tm of growth for the same ratio (XsinalX')=0.06. The dynamic range linearity (Fig.7) and errors (Fig.

The dynamic range linearity (Fig.7) and errors (Fig. 8) were estimated by placing a sample that was a flat surface of silicon carbide at the plane G = 0 and moving the sample in the Z direction.120- - • - Xsina1X'=1130. E 15.E .f!! c "tl .-8.Ul100- - Xsina1X'=111680---XsinaIX'=1/1040 L~~~~==='-:-:"':=-:'===-=. --60_ - ••..- .-- '"'"' ~20 o--'-.... ""._"""," ," ,o3500300010002000150050025004000Z scanning in urnFigure.7 The displacement of the spot (mm) function of the Z-scanningGrowth detection error (um)2 01 0o Z -5 can m m1 020Figure. 8 Growth detection error in urn420ConclusionThe above paper describes continuing progress toward a SALD technology capable of producing prototype as well as structural parts. The current process already yields dense silicon carbide free of the glassy binder phases which typically comprolnise high telnperature resistance.

The current process already yields dense silicon carbide free of the glassy binder phases which typically comprolnise high telnperature resistance. The alternative precursor, diethylsilane, shows incompatibility with CO2 lasers, but growth at low pressures with a greatly attenuated beam indicates its potential for lower deposition temperatures. The process continues to produce higher growth rates, and thermal shock and contamination of the optical path no longer pose significant obstacles in the practicallnatter of producing large scale parts. Additionally, an in-situ method for measuring height of the material following deposition represents a step toward actively controlling the unifonnity of deposition height in spite of variations in deposition rate.

Deposition Vapor Infiltration" , Proceedings of the Solid Freeform Fabrication Symposium, edited by Harris L. Marcus, 1. 1. Bemnan, David L. Bourell, and 1. W. Barlow, The University of Texas at Austin, Austin, Texas, August 8-10, 1994.Birminghmn, B. R. and Harris L. Marcus, " Silicon Carbide Shapes by Selective Area Laser421.

VANCES IN SLS POWDER CHARACTERIZATIONA. Amado*, M. Schmid†, G. Levy† and K. Wegener*Department of Mechanical and Process Engineering, Swiss Institute of Technology, Zurich 8008, Switzerland †inspire AG, irpd Institute, St. Gallen 9014, SwitzerlandAbstractThis paper introduces a new Selective Laser Sintering (SLS) powder characterization methodology. A better understanding regarding powder flow processing range is targeted. Intrinsic properties of polymers are given from the basic chemical structure and non-intrinsic ones describe characteristics caused from pre-processing and production. The non-intrinsic ones are dedicated to the powder. Understanding the particle size distribution and shape coupled with its ability to flow under the particular SLS processing conditions is desired. In this direction, a system called Revolution Powder Analyzer is employed and the dynamic powder behavior is characterized in nearly roll spreading conditions.

In this direction, a system called Revolution Powder Analyzer is employed and the dynamic powder behavior is characterized in nearly roll spreading conditions. This allows a sensitive differentiation of powders regarding their flow-ability and predicts, to a certain extent, the behavior under SLS conditions.1 Introduction1.1 SLS Powder Development: State of the artThe development of new powdered materials suitable for the Selective Laser Sintering Process with the aim to broaden its application field constitutes one of the main research topics and challenges nowadays [1,2,3,4]. The types of materials actually used are mostly neat polymers, particularly polyamides or polyamide based compounds [5], achieving a market share of 95% [3]. Several research efforts have being conducted during the last years towards the design of new SLS powders (see Table 1).

Several research efforts have being conducted during the last years towards the design of new SLS powders (see Table 1). In this direction, different approaches have been adopted, considering just from a simple mixture of two polymer components to more sophisticated production methods like mechano-chemical alloying [6,8]. Depending on the production method employed, intrinsic property changes can be induced, such as a shift of the melting or crystallization point. However, non-intrinsic features like the particles size and shape or powder flowability are strongly the materials characterization for its SLS process suitability, most researchers concentrate their effort on intrinsic properties (Table 1). Thermal and rheological measurements conducted by differential scanning calorimetry (DSC), termo-gravimetric analysis (TG), melt flow index evaluation (MFI) and rheometer testing methods (viscosity characterization) constitute a standard practice.

However, non-intrinsic properties are normally barely reported or simple left out. It is well known that good dispersion conditions are necessary to achieve a higher powder packing and a homogeneous layer spreading, but no quantitative information towards a prediction for SLS powders is available in literature. Most researchers conduct a powder development cycle just by trial and error carried out on a full or scale SLS equipment. Salmoria et al. [7] investigated the processing of blends of commercial polyamide 12 (PA12) and high density polyethylene (HDPE) powders. A year later, a similar approach was performed, considering blends of polyamide 6influencedin a sensitive way [9]. Regarding(PA6) and PA12 [10]. Sintered parts with high porosity were obtained and a powder pre- treatment is mentioned to improve the particle regularity.

Regarding(PA6) and PA12 [10]. Sintered parts with high porosity were obtained and a powder pre- treatment is mentioned to improve the particle regularity. However, no clear information regarding particle size distribution (PSD) or shape was reported, which does not clearly allow distinguishing between a predominant intrinsic or non-intrinsic effect. Schultz et al. [8] investigated PA12-PEEK blends formed by cryogenic mechanical alloying. In this case, due the processing method employed, a flake-like structure of the alloyed particle is reported and a large amount of fines are revealed by the PSD analysis. The author attributes the lower mechanical properties to the difficulties in achieving a dense powder bed despite having achieved a good miscibility between both materials. More recently, the modification of polyamide with organic and inorganic nano-fillers to enhance polymer performance, i.e., fire retardancy, high strength and high heat resistance has been employed. Koo et al.

Koo et al. [11] examined polyamide 11 and 12 modified with different types of nanoparticles produced by a twin extruder followed by a cryogenically grinding step. The subsequent trial and error cycle proved that not all blends were successful. The author states that the reason for the failure was the process inhibition by powder mechanics. The poor powder flow led to poor powder deposition and subsequent SLS processing difficulties.

Even the rejected material is described as flour like flowing powder, but again a qualitative appraisal was elucidated.Production TechnologyPowder mixing (polymer, fiber, beads)Melt mixing & cryogenic grinding/spray dryingDissolution- precipitationMechano- chemical alloying/ Solid stateIntrinsic characterization testAuthor/ Research GroupNon-intrinsic characterization testParticle Shape         [7][10][12]         [13][14][15][16]                                 [11][20]                                                MFI/ Rheo.Tap/Bulk DensityDSC**OthersTGOthers PSD[17][18][19][21][22][23][6][8][9][24]Other characterization methods: *FTIR, EDX, XRD, etc.; **Angle Of Repose, Carr Index, etc.

Table 1: Material characterization methods employed by different authorsTable 1 summarizes for different authors/research groups the diverse production technologies employed and the materials characterization methods used prior and after carrying out the sintering trials. As noted, most of them did not consider any kind of powder flowability evaluation despite that each production process has a considerable different effect on the non- intrinsic powder properties. Normally just the particle shape characteristic is reported as irregular. Actually, there is no quantitative measurement technique in the existing SLS field to characterize the powder flow behaviour. This makes it very difficult to predict the process-ability of a new material a priori.

This makes it very difficult to predict the process-ability of a new material a priori. Thus, it is suggested that a quantitative powder mechanics characterization methodology must be developed to facilitate the SLS processing [11].1.2 Non-intrinsic powder characterizationThe flow properties depend on many parameters, e.g., particle size distribution, particle shape, inter-particle forces, moisture and temperature [25]. It is a challenging task to determine theoretically the flow behaviour of bulk solids in dependence of all of these parameters. Thus it is necessary to determine the flow properties in appropriate testing devices. Evans et al. [26] introduced a “SLS Materials Development Method” were he emphasizes the importance of considering the powder behaviour as a first step on a powder development cycle. However, currently no specific method has been established. Many characterization techniques are available to determine the flow properties of powders. Krantz et al.

However, currently no specific method has been established. Many characterization techniques are available to determine the flow properties of powders. Krantz et al. [27] provide a comprehensive description of different techniques, considering static and dynamic powder state conditions.

Krantz et al. [27] provide a comprehensive description of different techniques, considering static and dynamic powder state conditions. Table 2 gives an overview of the most common techniques in use, particularly in the food and pharmaceutical industry.Bed Expansion RatioAngle of ReposeRing Shear CellBulk/Tap DensityMethodMeasurement ConditionCharacterization parametersStandardDynamic under vertical fluid drag load Fluidized height v/s upstream fluid flow Not standardizedStatic under free external loadPile angleDIN ISO 4324Quasi-Static under pressureShear force v/s normal pressure/ compression rate ASTM D6773Static under the effect of powder weight Loose and packed height v/s n° of taps ASTM D7481Table 2: Most common powder characterization methods [27]It was established that results provided by each method are strongly dependent upon the powder stress condition.

Therefore, techniques that consider an aerated state as the bed expansion ratio are appropriate to predict the fluidization performance, while externally loaded powder methods give an indication about static stability and compressibility ratio. Thus, no single technique is suitable for the full characterization of a powder and all of them, in principle, complement themselves. While more information is known about the handling system, the more likely it is to choose an accurate measurement device. Here must be emphasized to select a suitable characterization method that can replicate as closely as possible the boundary condition requirements. Unfortunately this is possible just for simple powder handling systems and geometries.

Unfortunately this is possible just for simple powder handling systems and geometries. For the SLS spreading system, a rather difficult appraisal about the deposition conditions can be assessed since no clear information is available to estimate if the stress state is next to a fluidized or a more compacted stress state.2 ObjectiveThe aim of this paper is to introduce a new powder characterization method into the SLS field that gives information regarding the dynamic behavior under a similar stress state when powder is spread over the part bed. The intent is not to replace traditional methods, but complement those existent to achieve a more accurate understanding about SLS powder suitability and thus reduce the powder development cycle time.3 Experimental Device & Methodology3.1 Dynamic Powder System: Operational PrincipleThe measurement device employed consists of a Revolution Powder Analyzer manufactured by Mercury Scientific Inc.

It consists of a rotating and an image acquisition system as shown in Figure 1. The rotating drum is machined in aluminum with an inner diameter of 50 mm and 35 mm width. The lateral sides are covered with transparent glass, to allow the powder behavior inside be captured by the image acquisition system. The drum can be rotated at different speeds ranging from 0 to 200 rpm.DrumBacklightDrum Cross Sectional ViewCameraPowder surface boundaryRollersFigure 1: Schematic diagram of rotating and image acquisition systemWith the aid of a backlight source, the powder free surface and cross sectional area of powder inside the drum can be recorded. Depending on the turning speed, two different tests can be performed: At low values, a discrete behavior is achieved based on a sequence of avalanches; at higher speeds, a continuous operation mode is reached, characterized by a steady state regime. The tested behavior is called Flowability and Fluidization respectively.

The tested behavior is called Flowability and Fluidization respectively. For the present research, the initial set up and the different parameters employed for each method are summarized in the following table:Flowability TestFluidization TestParameter Sample Volume Rotating Speed Preparation Time Avalanche Threshold Angle Calculation N° Avalanches to record Image capturing rateValue 25 cc (tap density) 0,6 rpm 30 s 0,65 % Half 128 15 fpsParameter Sample Volume Prep.

Rotating Speed Preparation Time Initial Rotating Speed Final Rotating Speed Rotating speed increment Image capturing rateValue 25 cc (tap density) 90 rpm 30 s 50 rpm 90 rpm 10 rpm 30 fpsInternal drum surface roughness: Ra=2.6 m Table 3: Flowability and Fluidization test set up main parameters (RH: 40%; Room T°: 25 °C)It must be pointed out that the drum rotational speed range for the Fluidization test can be adjusted to similar angular speeds of the translating roller during the powder deposition cycle (e.g., 35 to 80 rpm at a translational speed of 77 to 177 mm/s for a 3DSystems Sinterstation 2000 machine equipped with a 50 mm roller diameter).It is important to remark that with this method a nearly similar stress state condition can be achieved in comparison to the techniques presented previously (Table 2).

The boundary conditions generated for the powder surface inside the drum permit emulate the typical front stress free turning powder wedge behaviour generated by any of the actual SLS spreading systems, i.e., the counter-clockwise rotating roller (3DSystems) or the concave blade coater (EOS) (Figure 2).Figure 2: SLS powder spreading systems: Roller (3DSystems) & Blade (EOS) [28]3.2 Characterization IndexesBased on the sequence of images captured by the acquisition system, the following characterization parameters for each image are developed according to the specific test performed:Avalanche Angle: corresponds to the angle obtained from a linear regression of the free surface at the maximum potential energy prior to the start of the powder avalanche occurrence. Normally the left half of the diameter is considered to obtain a more representative measure. As a general rule, the higher the avalanche angle the poorer the flowability.

Normally the left half of the diameter is considered to obtain a more representative measure. As a general rule, the higher the avalanche angle the poorer the flowability. In comparison to the angle of repose presented previously in Table 2 (DIN ISO 4324), a distribution of values is obtained.Surface Fractal: corresponds to the fractal dimension D of the free surface of the powder and provides an indication of how rough the powder surface is. The determination of this dimensionless parameter is based on the method used by Richardson [29] who proposed the empirical relationship relating the length estimate, L() ,with the scale, , of measurement given by:L() = M (1-D)where M is a positive value and D is a constant at least equal to unity.

In our study the length estimate is related to the powder surface length and the scale is varied between a minimum limit that is defined by the image resolution (i.e., the pixel size) and a maximum that normally corresponds to one third of the drum diameter.Flowability Test Avalanche Angle Surface Fractal Volume Expansion RatioFluidization Test Fluidized Volume Slope Fluidized Height Slope Final Settling TimeAva. AngleFluid. VolumeSurface FractalFluid. WidthLow SpeedExp. VolumeHigh SpeedFluid. HeightTable 4: Flowability & Fluidization IndexesThe measurement is performed after each avalanche to determine how the powder reorganizes itself. If the powder forms a smooth even surface, the surface fractal will be near one.

If the powder forms a smooth even surface, the surface fractal will be near one. If the surface is rough and jagged, the surface fractal will be greater than one.Volume Expansion Ratio (VER): corresponds to the ratio between the volume measured inside the drum (expanded volume that relates to the bulk density) and the volume occupied by the powder in the sample container before filling the drum during the preparation stage (tap density). The preparation step consists of filling a 25 cc cylinder under manual tapping until the maximum powder compaction is achieved. Then the upper surface of the cylinder is scraped out to accurately define a fixed volume for all samples.

Then the upper surface of the cylinder is scraped out to accurately define a fixed volume for all samples. The expanded volume is measured as the sum of the area of every image pixel occupied by the powder multiplied by the width of the drum.Fluidized Volume Slope: is defined as a linear regression analysis of the cumulative volume that is fluidized versus the angular velocity of the sample during the fluidization testing. The fluidized volume corresponds to the cross sectional area inside the drum which develops a near horizontal powder surface.

The fluidized volume corresponds to the cross sectional area inside the drum which develops a near horizontal powder surface. The limit of this area is set when a high steep change of the free surface slope is generated (see Table 4).Fluidized Height Slope: is defined as a linear regression analysis of the cumulative height of the fluidized volume versus the angular velocity of the sample during the fluidization testing.Final Settling Time: corresponds to the time elapsed until the sample stops settling from the moment the drum rotation is stopped at the maximum rotational rate.

In this case the powder settling corresponds to the reduction of the fluidized height versus time.Also, for each powder tested, 3 consecutive measurements were considered, to determine any possible variations regarding the repeatability of the results.Among the parameters presented, the ratio between the volume expanded and the sample volume (VER) can be correlated to a certain extent to the so called Hausner Ratio (HR) or Carr index (CI=1-HR-1) (ASTM D7481) derived from compressibility studies. Both indexes are well accepted due to their relative ease of determination and excellent reproducibility in comparison to other indicators like the angle of repose (DIN ISO 4324). However, regardless of their wide applicability, they present some drawbacks [30]. Usually similar materials are not likely to be well differentiated. In fact, two materials with different bulk and tap densities but similar HR values are likely to behave very differently in practice [31].

In fact, two materials with different bulk and tap densities but similar HR values are likely to behave very differently in practice [31]. Thus, a deeper understanding and adaptation towards specific applications like SLS are needed.4 Results & AnalysisA selection of different SLS commercial polymer based materials and others under current research were tested with the aim to characterize their properties in relation to the indexes presented in the previous chapter. Powders under development, due a non-disclosure agreement, are labelled as Materials 1, 2A and 2B. The difference between A and B is a post-processing Spray Drying treatment to improve the particle shape (B).

The difference between A and B is a post-processing Spray Drying treatment to improve the particle shape (B). Table 5 provides a summary of the materials chosen and their principal features, including a PSD analysis on a volume basis considering an equivalent circular diameter.MaterialDF PA12PA 2200AlumideDF HST DF FlexicoPPMaterial 1Material 2ABase MaterialPA 12PA 12PA 12PA 12TPE*coPPPAPAFillerNoneNoneAluminiumMineral FibersN/ANoneNoneNano- SilicaCompound MethodNeat PolymerNeat PolymerMech. MixingMech.

MixingMech. MixingNeat MaterialNeat PolymerNeat PolymerMechano FusionFinal Compound1 Phase1 Phase2 Phases2 Phases1 Phase1 Phase1 Phase1 PhasePSD (m)D10=44 D50=56 D90=76D10=39 D50=53 D90=72D10=40 D50=65 D90=85Not defined**D10=64 D50=88 D90=98D10=36 D50=58 D90=88D10=40 D50=48 D90=65D10=45 D50=97 D90=115Thermoplastic Elastomer; ** Not representative results for PSD with high different aspect ratios Table 5: SLS material properties tested with Revolution Powder AnalyzerFor each material both proposed tests were performed considering the configuration previously described in Table 3. This set of materials was selected because it represents diverse compounding and mixing methods to obtain powders with heterogeneous and homogeneous mixed phases. Additionally PA12 based materials constitute at least 90% of the market share.

Additionally PA12 based materials constitute at least 90% of the market share. Figure 3 depicts a sample image of each powder tested (same scale for all pictures).Figure 3: Image sample of SLS powders testedThe Flowability and Fluidization results are presented as follows:Material 2BPANano SilicaSpray Drying1 PhaseD10=22 D50=44 D90=704.1 Flowability Resultss e h c n a a v alla t o t18 , 0DF PA12 PA2200 Alumide DF HST DF Flex icoPP Material 1 Material 2A Material 2BLow Spanf o n o i t c a r FHigh Span80Avalanche Angle (degrees)Figure 4: Cumulative Avalanche Angle DistributionFigure 4 depicts the cumulative avalanche angle distribution for each material. As noted, different materials present different median values and also different curve symmetries (skewness). A low median angle coupled with a narrow value distribution (span) is correlated to a good flowing powder.

A low median angle coupled with a narrow value distribution (span) is correlated to a good flowing powder. In this case, icoPP, Material 1 and Material 2B present the lowest average D50 of 38°. These three powders present in common a fine PSD with a very near spherical shape and the presence of 1 phase. Following this characterization, DF PA12, PA2200 and Alumide constitute a second group with similar results. The average avalanche angle D50 achieves a value of 44° and these powders present a higher distribution of shapes with convex structures on its surface in comparison to the previous group (no relevant differences between 1 or 2 material phases). Finally, particles with a less geometrical defined shape or compounds with different shape aspect ratios, i.e., fibers or more complicated geometries like flakes can be grouped in a set of materials that present the highest avalanche angle and broader curve distributions.

In this case, a trend is difficult to assess due the heterogeneity of the materials (DF HST, DF Flex and Material 2A).0 0 1s e h c n a a v a f o r e b m u Nl0 9DF PA12 PA2200 Alumide DF HST DF Flex icoPP Material 1 Material 2A Material 2B9Surface Fractal DFigure 5: Cumulative Surface Fractal DistributionThe surface fractal results are depicted in Figure 5. In this case, in comparison to the avalanche distribution behaviour, this index does not distinguish between near spherical or broader shape distribution based powders, which can be grouped together with an average fractal dimension D50 of 1.72 (DF PA12, PA2200, icoPP, Material 1 and Material 2B). However, a second set of powders constituted by compounds with two heterogeneous phases, namely a polymeric matrix and filler, i.e., Alumide and HST, can be clearly differentiated with an average D50 value of 2.3.

Concerning a correlation to the SLS spreading process, these results indicate that the free boundary generated after the avalanche for this second group presents a more structured or jagged surface that can influence the packing of the powder. Extreme distorted particle geometries like of Material 2A and DF Flex present clearly higher fractal values and thus, a lower packing.Figure 6: Summary of results for Flowability IndexesFigure 6 summarises the results presented above for the median value of each index. Additionally, the volume expansion ratio (VER) is included and the bulk density of each powder measured inside the rotating drum. As observed, two main trends can be established. First, as the avalanche angle increases, the VER increases as well. This is in accordance with other results described in the literature.

As observed, two main trends can be established. First, as the avalanche angle increases, the VER increases as well. This is in accordance with other results described in the literature. This VER indicator, due its definition, can be correlated to the so called Hauser Ratio described previously, which relates the tap and the bulk density of a powder. However, this index fails to describe accurately the behaviour of DF Flex. Concerning the VER (HR) classification, DF PA12 and DF Flex should present the same ability to flow, which disagrees with the empirical evidence (optimal processing conditions on 3DSystems equipment are completely different from DF PA12 to avoid streaking on the powder bed). Second, in general, an opposed effect between bulk density and avalanche angle can be observed. As the bulk density decreases, the avalanche angle increases.

Second, in general, an opposed effect between bulk density and avalanche angle can be observed. As the bulk density decreases, the avalanche angle increases. In case of Material 2A, despite having a lower density and higher VER in comparison to DF Flex, it presents a considerable lower surface fractal dimension. Therefore, both trends do not allow ranking adequately by themselves the powders under study in relation to a proper packing performance. Thus, based on the resultspresented, a new indicator is proposed, based on the ratio between the VER and the Surface Fractal Dimension, i.e.: Rank Index=VER/D As depicted in Figure 6, this index can better rank the powders regarding its SLS suitability. The higher score is obtained for Material 2B (spray dried) and the two commercial PA12 in agreement with the empirical experience on SLS equipment and their ease of processing.

The higher score is obtained for Material 2B (spray dried) and the two commercial PA12 in agreement with the empirical experience on SLS equipment and their ease of processing. Materials icoPP and Material 1 can be classified below the neat polyamides and finally, both filled materials with a heterogeneous phase, i.e. Alumide and DF HST follow in this classification. Finally, DF Flex results to be the lowest ranked material followed by the flake like powder Material 2A.

In general, the author propose that this index constitutes an improved description due its ability to couple information regarding the pile stability and powder rearrangement, which can better explain how DF PA12 and PA2200 fulfil an adequate compromise between these two factors.4.2 Fluidization ResultsRegarding the test at higher rotational speeds, the results for the fluidized powder height and volume are depicted in Figure 7 & 8.)m c (t h g e H d e z d u F e v i t a u m u CiiillDF PA12 PA2200 Alumide DF HST DF Flex icoPP Material 1 Material 2A Material 2BDrum speed (rpm) Figure 7: Cumulative Fluidized Height) c c ( e m u o V d e z d u F e v i t a u m u CliillDF PA12 PA2200 Alumide DF HST DF Flex icoPP Material 1 Material 2A Material 2BDrum speed (rpm) Figure 8: Cumulative Fluidized VolumeBoth figures depict a linear correlation between the drum speed from 50 to 90 rpm and the fluidized height and volume.

As noted, materials with a finer PSD and near spherical shape present the maximum volume slopes. For the fluidized height the same trend is observed. It must be pointed out that the term “fluidized” refers specifically to a powder that presents a bulk volume expansion due the entrapped air or gas during powder agitation when the fluid evacuates the powder interstices at a lower rate than the fluid that is introduced inside the inter-particle voids. However, this condition might not be necessary to achieve a volume expansion. Thus, it must be distinguished between the effect of a real fluidization or just a volume dilatation due the reduction of the coordination number when particles move in a shear displacement relative to each other. This phenomenon is known as dilatancy1 [32]. This effect can be analysed by the Final Settling Time parameter.

This phenomenon is known as dilatancy1 [32]. This effect can be analysed by the Final Settling Time parameter. If after the drum detention this value turns out to be negligible, it indicates that the particles remain in contact due their interconnections and the powder does not really fluidize.Figure 9: Summary of results for Fluidization IndexesAnalysing the results depicted in Figure 9, it can be observed that the first three materials present the highest settling times coupled with the maximum height and volume slopes. It can be stated that these powders develop a fluidized state. DF PA12 and PA 2200 depict lower settling time values correlated to an incipient fluidization. On the other hand, both filled materials with heterogeneous phases (Alumide & DF HST) and DF Flex clearly show a dilatancy effect with lower volume expansion rates. DF Flex seems even to be invariant to the drum rotational speed.

DF Flex seems even to be invariant to the drum rotational speed. This can explain why it presents a lower VER value as the bulk volume remains almost unchanged from its compacted condition (tap density). In case of Material 2A a particular condition occurs. The powder partially fluidizes with a lower volume expansion rate and a higher1 For the Fluidization test its name is used indistinctively for any of both conditions, i.e., a real fluidization or dilatancyheight slope. From the flowability results, this powder presents the maximum VER value. This indicates that the powder easily reaches a low bulk density (high expanded volume) from its initial compacted condition (tap density) under slow rotational speeds. That can be achieved only due a dilatancy effect.

That can be achieved only due a dilatancy effect. Therefore, this material, due to its particle flake shape, increases its height just at the top free surface due the avalanching effect coupled to a high interlocking between particles, but without increasing the fluidized volume width2, which remains quite similar to DF PA12 (see Table 4).Concerning the SLS powder deposition, a distinction between fluidization and dilatancy could predict to a certain extent how the deposition speed can affect the packing conditions. If the powder presents a higher fluidization rate like Material 2B, a higher speed will lead to a poor packing state. Thus, a lower translational velocity of the spreading system must be considered. On the other hand, if the powder density condition is less sensitive to this variable, a relative constant packing will be obtained.

On the other hand, if the powder density condition is less sensitive to this variable, a relative constant packing will be obtained. However, if a strong dilatancy effect is present, like for Material 2A, a lower part bed density can also be obtained.5 SummaryA new SLS powder characterization methodology has been introduced considering the dynamic tumbling behavior under a similar processing stress state when the powder is mechanically agitated inside a rotational drum. The Revolution Powder Analyzer system was employed and different characterization indexes were described and correlated to the SLS process.Different commercial powders and others under research were analyzed. New aspects regarding powder rearrangement were coupled to traditional measurement indexes to improve the description of the powder packing performance.

New aspects regarding powder rearrangement were coupled to traditional measurement indexes to improve the description of the powder packing performance. Thus, a new characterization indicator is proposed, which could help to rank more accurately newly developed powders, assessing a sensitive differentiation.The intent is not to replace traditional methods, but complement those existent to achieve a more accurate understanding about SLS powder suitability and thus reduce the powder development cycle time.Of course, it must be remarked, that an adequate powder spreading and packing is just one of the fundamental conditions that must be met to successfully sinter a material and even a good flowing powder that presents unsuitable intrinsic properties can be rejected.Finally, this research was limited to new materials.

CURACYAUGUST 1991Dr. pauJ..iF. J'acobs Dire.ctor of Process TechnologyandMr. JanR.ichter Process Engineer3D Systems Inc. 26081 Avenue HallValencia, California91355138INTRODUCTIONIt has been almost four years since the SLA - 1 ushered in the new technology of StereoLithography, and about 2\ years since 3D Systems introduced the SLA-250. Since then, nearly 300 systems have been installed worldwide and are currently providing benefits in a range of applications which might well be summarized by the term "Rapid Prototyping and Manufacturing" or "RPM".the accuracy of parts built with stereoLithography has benefitted significantly from nine important technological advances. The research and development efforts which formed the foundation for this progress originated within the Process, Chemistry and Software departments of 3D Systems.During the past yearThe following is a listing, and brief description, of the keyfeatures of each of these advances.1.

WEAVETMThis proprietary building method was developed at 3D Systems during 1990. WEAVE~ is a trademark of 3D Systems, and a patent application for this technique has been filed. The WEAVEN building shrinkage method SUbstantially reduces part swelling, uniformity, dimensional instability (creep), enhances "green strength" and significantly is currently operational on improves overall part accuracy. WEAVE N the SLA-190, SLA-250, and SLA-500.improves anddecreasespost-curedistortion2. SOFTWARE RELEASE 3.81 with IMPROVED RESIN LEVELINGthe leveling operation was performed shortly after the z stage / platform had moved downward during the recoating sequence. the platform resulted in resin free surface oscillations which were not completely damped when level sensor sampling occurred. The result was long leveling intervals, involving mUltiple plunger corrections which themselves caused additional surface perturbations reSUlting in errors as great as ± the resin 2 mils.

The result was long leveling intervals, involving mUltiple plunger corrections which themselves caused additional surface perturbations reSUlting in errors as great as ± the resin 2 mils. with release 3.81 for the SLA-190 and SLA-250, level is now sensed after laser drawing. Consequently, neither the platform nor the leveling system plunger have moved for the longest possible time, and thus the fluid surface has very nearly reached stable equilibrium. the leveling system is now result, operating with a standard deviation of about 0.2 mil and a maximum error of ± 0.5 mil.Formerly,The motion ofAsa1393. NEW POST-CURING APPARATUS (PCA)The earlier PCA utilized compact, high intensity mercury the development of lamps. The new thermal stresses and considerable post-cure distortion.

NEW POST-CURING APPARATUS (PCA)The earlier PCA utilized compact, high intensity mercury the development of lamps. The new thermal stresses and considerable post-cure distortion. lamps, provides a nearly optimal PCA utilizes actinic fluorescent irradiance distribution resulting in more uniform post spectral cure, temperatures, greatly reduced thermal stresses, and improved part accuracy. The PCA-2S0 applies to all resins used in either the SLA-190 or the SLA-2S0. The PCA SOO is optimized for and also the resins used in the SLA-SOO, accommodates correspondingly larger parts.Test data showed excessive heating,significantly lower part4. TWIN-SCREW ADJUSTABLE RECOATER BLADEThe primary function of the recoater blade is the accurate and reproducible generation of uniform fresh resin layer thickness. Errors in blade clearance can lead to delamination, excessive curl, and reduced overall part accuracy.

Errors in blade clearance can lead to delamination, excessive curl, and reduced overall part accuracy. recoater blade time consuming process with a mean error adjustment was a tedious, of about ±·2 mils. with the implementation of the new apparatus, the time required to set the blade gap has been significantly reduced and the mean error in blade gap is now down to ± O.S mil. layer The The new thickness and reduced incidence of twin-screw. blade has been retrofit on earlier SLA-2S0' s and is currently installed on all new SLA-2S0's.Previously,resultis more accurate recoating, more consistentlayer delamination.5."WINDOWPANE" METHOD FOR DETERMINING RESIN PARAMETERS(penetration (critical exposure) were determined from "working depth) curves" (an experimentally determined plot of polymer cure depth the logarithm of the laser exposure) which exhibited standard vs. deviations in cure depth of O.S mil, and maximum variations of about ± 1.S mils.

deviations in cure depth of O.S mil, and maximum variations of about ± 1.S mils. Errors of this size could often negate all the • A new method, know by the visually potential advantages of WEAVEN descriptive title "WINDOWPANE", was designed, developed, tested, optimized, documented, and is now in use throughout 3D Systems. The new "WINDOWPANE" test results indicate standard deviations in cure depth of 0.2 mil, which is sUfficiently accurate to exploit , and is now the basis of quality control the advantages of WEAVE N for all for StereoLithography applications.Untilrecently,the key resin parameters Dpand Ecresinsproducedby Ciba-Geigy6. AUTOMATIC COMPUTATION OF EXPOSURE AND LASER DRAWING SPEEDWith accurate knowledge of Dp and Ec it is now possible to correctly determine the required exposure necessary to achieve a given cure depth.

Also, coupled with an accurate measurement of140the laser power at automatically computes required exposure and, ultimately,the resin surface,the current software now the proper drawing speed to achieve thethe correct cure depth.is no longer necessary for the user to manually insert step Cure depths generated with the current than thoseperiods or step sizes. software established by the prior method.are moreaccurateand moreconsistent1. MORE ACCURATE MACHINE CALIBRATION TECHNIQUEA new calibration technique was developed for the SLA-250. (approximately 3 hours vs. 13 The new method is not only fa.ter is also more accurate than the previous calibration hours), but Results were obtained using a coordinate measuring technique. machine long dimensions on a standard diagnostic test part specifically used for machine calibration.(CMM).Data wastaken forshort, medium and8. NEW PART CLEANING PROCEDUREstereoLithography parts were cleaned with alcohol.

NEW PART CLEANING PROCEDUREstereoLithography parts were cleaned with alcohol. showed that even a relatively brief exposure of "green" parts to alcohol resulted in significant part distortion (e.g. 8 mils of swelling on a 2 inch dimension of a part built with WEAVEN after exposure to alcohol for using the new part cleaning process, with TPM (Tri 0.5 hours). Propylene Glycol Monomethyl Ether) for uncured resin, the data shows dramatically reduced part distortion (e.g. less than 0.5 mil on the same 2 inch WEAVEN part dimension for exposures up to 5 hours!). Therefore, 3D strongly recommends that Detailed information new cleaning method. users regarding the requirements the new part for cleaning procedure will shortly be available from 3D Systems.Formerly, mostCareful experimental measurementsas the solventadopttheimplementation of9. STAR-WEAVETMadvanced building method was developed and tested.

STAR-WEAVETMadvanced building method was developed and tested. This proprietary technique, known as STAR WEAVEN, is also a trademark of 3D Systems, and a patent application for this technique has also been filed. STAR-WEAVEN extends the improvements already documented with WEAVEN through the addition of three new features. is derived from the inclusion of STaggered WEAVEN, Alternate sequencing, and Retracted Hatch.During1991aneven moreSpecifically,the acronym "STAR"STaggered WEAVEN involves offsetting the vectors of a given hatch pattern (e.g. the x vectors) on the (N+1)th layer, by exactly half a regUlar hatch spacing, relative to those on the Nth layer.141The result is a more closely Qacked density of cured hatch vectors.

the x vectors) on the (N+1)th layer, by exactly half a regUlar hatch spacing, relative to those on the Nth layer.141The result is a more closely Qacked density of cured hatch vectors. are also The vectors on the second pass staggered on consecutive layers.(e.g.the y vectors)Experiments have shown that STaggered WEAVEN results in a higher average level of cure for "green parts," increases average "green imQroves reduces swelling, and virtually eliminates dimensional stability. voids, drainage zones and micro-cracks. In general, STaggered WEAVEN results in parts with improved homogeneity.strength,"reducespost-curedistortion,Alternate sequencing reduces non-uniform part distortions as well as hatch dislocations tangent to interior openings, which are a consequence of using identical drawing sequences on each layer. The non-uniform distortions are quite probably due to layer-to layer accumulation of resulting from delayed shrinkage phenomena.

The non-uniform distortions are quite probably due to layer-to layer accumulation of resulting from delayed shrinkage phenomena. By alternating both the sequence and direction of propagation of the x and y hatch vectors drawn on successive layers, a partial cancellation of opposing moments can be achieved, with a corresponding reduction in part distortion. Also, hatch dislocations are no longer superimposed from layer-to layer cracks or extensive fissures. Software release 3.82 for the SLA-190 and SLA 250, in September 1991, will provide the capability to do STAR-WEAVEN with eight-fold Alternate eight possible combinations of two vector types with four propagation directions).internal bending momentsand hence do not cause structural weakness,and software release 1.4 forthe SLA-500,both duesequencing(viz.alternating allFinally, Retracted hatch involves attaching each hatch vector at only one border while retracting the hatch vector by a short distance (e.g.

20 mils) from the opposite border. Furthermore, the retraction sequence is itself alternated. ThUS, a given vector may be retracted from the left border while its adjacent neighboring In this way, vector would be retracted from the right border. the equal strength is preserved at all borders. As polymer cured on the first pass of STAR-WEAVEN, on any layer, is reaction forces. free to shrink without generating substantial bending Thus" Retracted hatch reduces moments. diminishes part distortion and provides a corresponding Test results for STAR-WEAVEN are improvement included herein.aresult,internalstressesandin part accuracy.142RESULTSregarding the to provide quantitative results accuracy of StereoLithography parts, it is necessary to select a particular accuracy test standard.

The ideal standard test part would have the following properties:In ordera.It would be large enough in x and y to insure that the system is capable of building accurately near the extremes of the platform as well as near the center.b.It would have a large number of small, medium, and large dimensions.c.It would dimensions. that linewidth compensation is working properly.

that linewidth compensation is working properly. Inside dimensions will tend to be too small, while outside dimensions will tend to be too large if linewidth compensation is either not operational or imprecise.haveboth"inside""outside" in order to checkandThis is importantd.It should not take too long to build.e.It should not consume a large quantity of resin.f.It should be easily measured with a CMM.g.It would have many of the features of "real" parts (e.g.thin walls, flat surfaces, holes, etc.).h.Ideally, Systems,it would be a part not designed by 3D to insure complete impartiality.Fortunately, such a part does exist. During the past year, about 150 the StereoLithography Users Group, industrial Government companies, laboratories, and Universities, developed exactly such an accuracy Known simply as the" User-Part", a photograph of test standard.

A drawing of the User-Part is this part shown in Fiqure 2, which also shows some of the many dimensions to be measured on a CMM. In all, a total of 170 measurements are made on each User-Part. Of these, 78 are in x, 78 are in y and 14 are in z.consisting ofservicebureaus,U.s.is shown in Fiqure 1,143(4\"."" "'J- M£A$UA.I.r\' ...,& ,,,,,..... FA.QM "l"""'~ "'1i.\(,""T)145FIGURE 2The proc;:edure iS Cl.$ folJ.ows:1.pick a build style. 60° 1120° Ix 50 mil recently, STAR-WEAVETII for this report, three of these methods.One might choose to build with . triangler or, In developing the data base 3D has built User-Parts with allhatch,WEAVE TIl2.Pick a resin. 5081-1, XB-5131, XB-5139 and XB-5143.3D has built the User-Part in XBpick compensation optimized.ashrinkagefaetora values willandvalue.Theselinewidth be later4.Build User~l?Clrt·.

this normally takes 8 overni.ght.theDePf3nding to 10 hours,uponpower and can be donelaser5.Clean the PCl:r:ti. done using TP~, which currently approved for StereoLithography.As noted earlier,this should be resinsis effective for all6.Postcure should using based PCA. Post-cure takes about 1 hour.the part. doneA9ad..n, theas noted earJ.i.er, new actinicthis fluorescentbe7.Measure the part. CMM.This takes about 20 minutes on a8.Fr.om the measured values of each of the 170 dimensions.thesemeasurements,determineactual9.Comparef3agh oft.he~Efactl.1al dimensi()ns with-the appropriate. CAD dimen$iion. which is the error ·for that speci.fic. measurement.Calculate the.

the the error the desiredof describesany17. Plot the 99% confidence limit (i.e. 3 standard deviation) error level as a function of scale. This plot describes how the errors are distributed for small, medium or large dimensions.18. Change the values of the shrinkage factor and linewidth compensation to reduce the errors.19. Repeatthe entire procedure, using the new shrinkagefactor and linewidth compensation value.20.continue achieved.iterating untiltheoptimum resultsareto note that once this process has been completed, for a particular resin and build style, all sUbsequent parts that are built with the same resin in that style, using the optimum shrink factor, level of should achieve the reason for independently determining the RMS accuracy. Also, errors in x and y, as well as the RMS error for all measurements, If the RMS errors for the x is to check the machine calibration. and y dimensions are very nearly equal, then the SLA is properly calibrated.

and y dimensions are very nearly equal, then the SLA is properly calibrated. However, if they differ significantly, then the system will probably benefit from re-calibration.Itis importantthe highestFigure 3 shows the error distribution function for a single in • A numberUser-Part built with resin XB-5081-1 on an SLA-250 (SIN 90-076) the Process Department at 3D Systems, using STAR-WEAVEN of important features of this data are worth noting:1.within the limitations of a finite sample (viz. N=170 data is nearly Gaussian.points),theerror distribution function2.the The function is also nearly symmetric (i.e. almost same number of positive errors as negative errors). The fact function is not perfectly symmetric suggests room for further improvement.thatthis1471 4 8F IG U R E3 ·. -. .;' - ' , , , - . ,_ " 0 -;, - .. - ..- :, . . - " ! ; .'; : . ~. . .: .. -. .,.". -.. .- ...- , ;. .. .' -: ..

About 65% of the dimensions are within ± 3 ails of their CAD value. About 85% of the dimensions are within ± 5 Hils of their CAD value.4.However, about 2% of the dimensions have errors between 9 and 11 mils.Figure 5 is a plot of the 99% confidence limit error as a function of the length of the dimension. This plot is generated by taking all lie within a that certain range of from 0 to 1 inch, 1 inch to 2 inches, etc.), and then calculating the 99% confidence limit error for this sub-set.the dimensions of lengthsthe User-Part(e.

g.Inspection of Figure 5 shows the following:1.The 99% confidence limit error increases with increased length.2.is length Howt.;ver, nonlinear! "RPM" accuracy based on "percentage" of length, or "mils per inch", both in of which imply a agreement with experimental results.theincreaseoferror withThus, characterizations oflinearfunction, are simply not3.The actual results suggest that the 99% confidence limit the error scales very nearly as Thus, for example, 9 inch dimensions will tend length. to have errors inch 3 times dimensions, not 9 times as great!the square-root ofaboutas great as1From these results, it should be evident that characterizing the "Accuracy" of StereoLithography is not as simple as referring to a particular number, be that a percentage or a maximum error. At this time we stongly believe that the best measure of accuracy1491 5 0F I G U R E4\, - : ' : : ' . ': : ' . : J. • . . .' \. i ;' . . . . . . . ..

6. This is a plot of the RMS error for all 170 dimensions of the User-Part, for resin XB-5081-1, when built with three different build methods.50 mil 60 0 /120 0 /x triangular hatch. This was by far the most common build method as little as one year ago.WEAVEN. This build method, in widespread use.intoduced late in 1990,STAR-WEAVEN• This latest build method will be available to the user base in September 1991.It should be noted that the data presented in Figure 6 are for parts generated after completing the iteration cycle a number of times. Thus, both linewidth and shrinkage compensation have not only been applied, but they have been optimized for each building method.Also, we have chosen to plot the results for the "3 best" User Parts built to date as a separate grouping.

This was done solely to illustrate the envelope of the present state-of-the-art.It is clear from the results presented in Figure 6 that WEAVEN in part building accuracy advance represented a relative to triangular hatch, and that STAR-WEAVEN represents an additional advance beyond WEAVEN•substantialin interpreting the data of ~iqur. 6 to recognize that if the error distribution function were truly Gaussian, the measurements then approximately 68% of all would lie within a tolerance band between plus and minus 1 RMS value relative to the CAD dimenion, about 90% within 1.6 RMS and about 95% within 2 RMS.Furthermore,it is importantFinally, it is not only important to characterize the accuracy of a single User-Part, but it also important to characterize the regeatability as well.

if one gets an accurate In other words, part on Monday, and keeps everything as constant as possible, what are the chances that one will still get another accurate part on Tuesday? To answer this question, a series of 15 User-Parts were all built with the same SLA, the same build technique and then cleaned, post-cured and measured as close tothe same resin,152is nowFigure 7 shows the error distribution identically as pO$sible. function for all 15 X.. 170 = 2550.meas\lremel'1ts. Figure 8 shows the cummulativeerror function for the same 2550 points. Notethat Figure 7 is very similar to Figure 3 and that Figure 8 is also·very similar to Figure 4.Analysis .of t):1edata shows that systemrep.eatability for the 15 user. Parts was within 1 milRMs\when usingSTAR-WEAVElM. This. is that onceproperbuild.methods and post excellent, proces$ing techniques are used, theSt.ereoLithography equipment is very repeatable.and show.s153~.< -. '--' ,.--.-., ...;.' ,... ~ ...,... . : ~ .,..

the User Part, StereoLithogra.Phy can now reliablymaintain70%>of all dimensions within ±5 mils, using STAR-WEAVEN.5.The good approximation, as the .sguare root of the •• length of a given dimension.99% confidencelimiterrorsscale,to6.System repeatability, based on a stUdy of 15 User Parts, is typically within 1 l1\ilRMS.IS7.

VANCES IN THE PRODUCTION OF INFILTRATED ALUMINIUM PARTS USING SELECTIVE LASER SINTERINGT.B. Sercombe and G.B. Schaffer Division of Materials, School of Engineering, The University of Queensland, St Lucia, Australia, 4072Reviewed, accepted August 20, 2004AbstractRecent advances in a rapid freeform fabrication process for the production of aluminium parts are considered. The methodology involves the formation of an unconstrained, resin bonded aluminium powder part by Selective Laser Sintering, the burnout of the resin, the partial transformation of the aluminium into a rigid aluminium nitride skeleton by reaction with the atmosphere under a magnesium/alumina blanket and the subsequent infiltration with a second aluminium alloy. Here we describe the process and consider potential applications.

Here we describe the process and consider potential applications. Strategies for controlling the growth of the aluminium nitride are also to be discussed.IntroductionThe automotive industry is both a major consumer of aluminium and the largest user of Rapid Prototyping (RP) technology (1). However, there is currently no commercialised technique to directly produce aluminium components using RP technology. In a recent paper (2), we described a new technique for producing such parts. It involves the production of a resin- bonded green part using Selective Laser Sintering (SLS), thermal decomposition of the binder, partial transformation of the aluminium particles into a percolating aluminium nitride skeleton by reaction of the aluminium powder with the nitrogen atmosphere, and subsequent infiltration of the preform with a second aluminium alloy. The formation of the skeleton is one of the key factors in the success of the process.

The formation of the skeleton is one of the key factors in the success of the process. It provides an inert rigid skeleton which is critical to maintaining dimensional stability of the preform during infiltration (3). However, the formation of such a skeleton is particularly problematical in aluminium because of the thermodynamic stability of the naturally occurring oxide layer that is present on the surface of the powder particles. The formation of AlN in a N2 atmosphere will only occur if the oxygen content is exceedingly low, <10-23 (4). This is not achievable in conventional atmospheres and therefore the use of an oxygen scavenger is required. Magnesium has a greater affinity for oxygen than Al, and therefore additions of Mg are made to the alumina support powder. This alumina powder provides support to the part for the period between removal of the binder and formation of the nitride skeleton.

This alumina powder provides support to the part for the period between removal of the binder and formation of the nitride skeleton. The Mg reduces the oxygen content below that where the oxide layer is stable, which then facilitates the formation of the AlN.The AlN skeleton also enhances infiltration by reducing the contact angle of the liquid. It has been reported (5) that the contact angle of molten aluminium on AlN is ~41° at 1100°C, approximately half that of aluminium on alumina. Therefore it is important that AlN forms throughout the entire preform. This is achieved by the addition of Mg to the support powder (6).The amount of nitride that forms is a critical parameter in the success of the process. Too little and dimensional stability is compromised, too much and cracking can occur (7). Additionally, rapid nitride growth rates makes the process extremely difficult to control, and80dramatically reduces the operational window, increasing manufacturing costs.

Additionally, rapid nitride growth rates makes the process extremely difficult to control, and80dramatically reduces the operational window, increasing manufacturing costs. In this paper, we describe the way in which we control the nitride thickness, thereby increasing the processing window and improving the mechanical properties of the resultant part.ExperimentalThe base powder used in this work consisted of pre-alloyed AA6061 powder that was mixed with 2% Mg and 4% (10v%) nylon. All compositions are listed in weight percent unless stated otherwise. To some parts, up to 2% Sn was also added. Selected properties of the powders used in this work are presented in Table 1. Parts were either produced in a SinterStation 2500plus (using a laser power of 32W, part bed and feed temperatures of 120°C and 70°C respectively and a layer thickness of 0.1mm) or heated to ~220°C for 30 min in a small steel mould.

This latter technique melted the nylon binder, which resulted in the formation of a resin-bonded green structure that simulated processing by SLS. The green preforms were placed in a crucible with a block of solid infiltrant at one end and covered with an Al2O3 – 1% Mg support blanket. Where only the formation of the skeleton was to be investigated, the infiltrant was omitted. The parts were then heated at 90°C/h in a sealed horizontal furnace to either 540°C or 570°C for up to 24h under flowing high purity N2.Table 1. Selected properties of the powders used.PowderCompositionParticle SizeAA6061 Al-1Mg-0.8Si-0.3Cu-0.1Cr53 + 10µmMg99.6% pure45µmSn99.8% pure45µmNylonco-polyamide 6-1210µmAl2O399.9% pure63µmInfiltration was effected by increasing the temperature at 90°C/h to either 570°C or 700°C for up to 7h. Infiltration was performed either under flowing N2 or Ar. If Ar was used, the furnace was purged for 1h at 540°C prior to further heating.

Infiltration was performed either under flowing N2 or Ar. If Ar was used, the furnace was purged for 1h at 540°C prior to further heating. Composition and melting ranges of the infiltrants used are given in Table 2. The melting ranges were determined using Thermo-Calc and the Al-data database (version 2, 2000) through the Ettan interface supplied by Thermotech Ltd.Samples for metallographic examination were mounted in epoxy and polished using standard techniques. They were left unetched and examined either optically or in a Phillips XL30 scanning electron microscope (SEM) with a LaB6 filament and fitted with an Oxford energy dispersive x-ray spectrometer (EDS) and backscatter detector.81Mechanical properties were determined using an Instron 1026 screw machine with a cross-head speed of 0.6mm/min. Strain was measured using an extensometer with a 50mm gauge length. The tensile bars had a parallel length of ~75mm and a cross sectional area of ~85mm2.Table 2.

Strain was measured using an extensometer with a 50mm gauge length. The tensile bars had a parallel length of ~75mm and a cross sectional area of ~85mm2.Table 2. The infiltrants used and their melting range.InfiltrantSolidus (oC) Liquidus (oC) Melting Range (oC)Al-14.3Si-4.7Mg Al Al-1Mg-1Mn Al-5Mg Al-2.4Zn Al-2.9Fe Al-2.2Zn Al-2.7Mn Al-2.3Cu Al-2.3Si557 660 641 580 650 654 639 658 604 577558 660 654 635 656 660 653 677 654 6471 0 13 55 6 6 14 19 50 70Results and DiscussionThe properties of 6061 - 2%Mg - 4%nylon preforms nitrided for 6h at 540°C and infiltrated with Al - 14.3%Si - 4.7%Mg for 4h at 570°C are presented in Table 3 in the as-cooled and solution treated condition. None of these parts exhibited any significant ductility. The fracture path, shown in Figure 1, indicates that the crack tended to follow the nitride. This suggests that the low ductility is a result of the percolating nitride skeleton providing an easy crack path.

This suggests that the low ductility is a result of the percolating nitride skeleton providing an easy crack path. In order to improve the properties, it is thus first necessary to control the nitride growth rate, which will allow more accurate formation of a given thickness of nitride layer.The addition of 2%Sn clearly reduces the rate at which the nitride grows (as measured by weight gain) at both 540°C and 570°C, Figure 2. Without Sn, the nitridation rate increases almost exponentially with time, while the addition of 2% Sn controls the growth to a much slower, linear rate. This reduced growth rate clearly produces a thinner nitride after a given hold time, as illustrated in Figure 3 for a hold of 6h at 540°C. The amount of Sn required to control the nitride is ~0.5wt%, Figure 4, with excess Sn having little effect for these conditions. The mechanism by which the Sn controls the nitride growth rate is not clear.

The mechanism by which the Sn controls the nitride growth rate is not clear. Kondoh et al (8, 9) suggest that the Sn forms a physical barrier, isolating the aluminium from the atmosphere. The fact that Sn has a low melting point and low surface tension tends to support this hypothesis. However, in the model proposed by Kondoh et al, the Sn formed islands on the particle’s surface, which would lead to non-uniform nitride thickness. The micrographs of Figure 3 do not support this. The mechanism by which the Sn controls the nitride is currently under investigation.82Table 3. Mechanical properties of selected alloys.

The micrographs of Figure 3 do not support this. The mechanism by which the Sn controls the nitride is currently under investigation.82Table 3. Mechanical properties of selected alloys. Properties of a common casting alloy is provided for reference.Base Powder 6061-2Mg- 4NylonInfiltrantAl-14.3Si- 4.7MgAl-14.3Si- 4.7MgAlNitride Time612 8 8 12Heat Treatment F T4@ F F 24h@540°C FYield (MPa) - -61UTS (MPa) 100 130 100 117 156 103Ductility (%) 0.4 0.5 0.5 0.5 0.6 2.06061-2Mg-1Sn- 4NylonAl-5Mg24 12 10 6Annealed Annealed Annealed Annealed121 86 82 77143 155 149 1420.6 1.6 2.5 36082Al-2Cu-1Mg12 12 612T4* T6* T6* F T4# T6#87 - 180 91 126 -118 184 209 124 143 2041.1 0.4 1.1 0.95 0.72 0.52Sand Cast 319.0-As cast T6^125 165185 2502 2F: as infiltrated; @ T4: 2h @ 500°C, water quench; * T4: 2h at 530°C, water quench. T6: T4 + 18h at 160°C; # T4: 2h at 540°C, water quench.

T6: T4 + 18h at 160°C; # T4: 2h at 540°C, water quench. T6: T4 + 18h at 200°C; Annealed: 2h at 410°C, water quench. ^ 12h @ 500°C, water quench, 2-5h @ 150°C.Figure 1. Fracture path for a 6061-2% Mg-4nylon preform infiltrated with Al-14.3% Si-4.7% Mg. This specimen was nitrided for 6 at 540°C and infiltrated at 570°C for 6h. The crack has tended to follow the nitride layer (arrowed) [from (2)].832016570oC No Sn540oC No Sn)%( n i a Gt h g i e W128570oC 2% Sn4540oC 2% Sn0036912Sintering Time (hrs)Figure 2. The effect of 2%Sn addition on the nitride growth rate (weight gain) of 6061 - 2%Mg - 4%nylon at 540°C and 570°C. The addition of Sn clearly reduces the rate at which the nitride forms.(a)(b)Figure 3. Microstructure of 6061 - 2%Mg - 4%nylon (a) with Sn and (b) without 2% Sn after 6h at 540°C.

The addition of Sn clearly reduces the rate at which the nitride forms.(a)(b)Figure 3. Microstructure of 6061 - 2%Mg - 4%nylon (a) with Sn and (b) without 2% Sn after 6h at 540°C. The addition of Sn has clearly resulted in a much thinner nitride layer [from (10)].84)%( n a Git h g e Wi36 32 28 24 20 16 12 8 4 000.511.522.5Sn Content (wt%)Figure 4. The effect of Sn content on the nitride growth of 6061-2% Mg-4% nylon after 12h at 540°C. As little as 0.5%Sn is sufficient to significantly reduce the nitride thickness.The addition of 1%Sn also increases the infiltration rate, as shown Figure 5.

As little as 0.5%Sn is sufficient to significantly reduce the nitride thickness.The addition of 1%Sn also increases the infiltration rate, as shown Figure 5. These curves were produced by removing parts from the furnace after various times at 570°C, removing and weighing the remaining infiltrant, and converting this to a percentage infiltrated, assuming that the part is 100% infiltrated when all the infiltrant has been consumed.10080d e t a r t l i f nI6040with SnNo Sn%2000120240360480Time at 570oCFigure 5. The effect of hold time on the infiltration rate of 6061 - 2%Mg - 4%nylon preforms with and without 1%Sn. The presence of Sn clearly increases the rate of infiltration. There is also a rapid increase in the infiltration rate of the Sn-containing alloys after ~15% has infiltrated [from (10)].Figure 5 shows that without Sn, there is an incubation period of ~2h, prior to the onset of infiltration.

The rate of infiltration (slope of the line) then increases and remains fairly constant, and the part is fully infiltrated after ~5h. In contrast, only a very short incubation period exists in85the presence of Sn, with noticeable infiltration occurring after as little as 30 minutes. After approximately 15% has been infiltrated, the rate increases sharply, producing a large amount of scatter at 90 minutes. Full infiltration has occurred after 2h. The enhanced infiltration may be attributed to the segregation of the Sn to the liquid interface, as illustrated in Figure 6. This figure shows a section through the advancing liquid, and the corresponding backscattered electron images from the liquid front as well as from the fully infiltrated zone behind the front. It shows that the infiltrated area has become depleted in Sn (the white phase), which appears to be concentrated at the advancing liquid interface.Figure 6. An optical cross section through a partially infiltrated bar.

An optical cross section through a partially infiltrated bar. The bar is 100 mm in length. The backscattered SEM micrographs are from the areas indicated. The Sn (white contrast) is concentrated predominantly at the advancing liquid interface.Tin has a low surface tension and is known to improve the wetting characteristics of a number of pressed and sintered aluminium alloys (11-13) . It is possible that Sn is having a86similar effect during infiltration. Once a critical Sn level has been reached in the advancing melt, wetting will be enhanced and rapid infiltration will follow. The result of this enhanced infiltration is that the time to fully infiltrate is halved.Despite the control of the nitride thickness due to the addition of Sn, the ductility of alloys infiltrated with the Al - 14.3%Si - 4.7%Mg eutectic remains low, Table 3. Even the thinner nitride formed after a shorter hold at 540°C (8h as opposed to 12h) had minimal effect on properties.

Even the thinner nitride formed after a shorter hold at 540°C (8h as opposed to 12h) had minimal effect on properties. A long term heat treatment aimed at spheroidising the Si improved the strength, but not the ductility. Investigation of the crack path, Figure 7, shows that failure is accompanied by cracking of the primary Si particles (arrowed). Thus it seems that the inherent brittleness of the Al-Si-Mg infiltrant causes premature failure of the infiltrated parts.Figure 7. Fracture path for a 6061 – 2%Mg – 1%Sn – 4%nylon preform infiltrated with Al – 14.7%Si – 4%Mg. This specimen was nitrided 8h at 540°C, infiltrated at 570°C for 6h, cooled to room temperature and then spherodised for 24h at 540°C.

This specimen was nitrided 8h at 540°C, infiltrated at 570°C for 6h, cooled to room temperature and then spherodised for 24h at 540°C. The crack has tended to follow the nitride layer, and is accompanied by cracking of the primary Si particle (arrowed).The usual requirement (3) of selecting an infiltrant with a melting point below that of the base powders (~580°C in this case) means that only highly alloyed infiltrants, based on the Al- Si-Mg ternary eutectic were able to be used. Since a percolating AlN skeleton has formed and AlN has a melting point in excess of 2000°C, the requirement for selecting an infiltrant with a melting point below that of the base powders is no longer applicable. However, since the nitride reaction will continue to occur until the infiltrant has covered the particles, variable nitride thickness is likely to result, with thicker nitride occurring in regions furthest from the infiltrant site.

This problem is exacerbated by the fact that the nitridation rate will be accelerated at the higher infiltration temperature. However, the nitridation reaction can be halted if the process gas is switched to Ar at the end of the 540°C hold. The difference in nitride thickness for 6061 – 2%Mg – 1%Sn – 4%nylon specimens that have been infiltrated in either N2 or Ar, thickness is illustrated in Figure 8. Both samples were nitrided for 12h at 540°C in N2. For the alloy shown in Figure 8(b), the process gas was then switched to Ar and the parts infiltrated for 6h at 570°C. Clearly, the sample infiltrated in Ar has resulted in a thinner nitride layer.87(a)(b)Figure 8. Microstructure of the 6061 – 2%Mg – 1%Sn – 4%nylon preforms infiltrated with Al – 14.3%Si – 4.7%Mg. Parts were nitride for 12h at 540°C and then infiltrated for 6h at 570°C in either (a) N2 or (b) Ar.

Parts were nitride for 12h at 540°C and then infiltrated for 6h at 570°C in either (a) N2 or (b) Ar. Clearly the nitride skeleton is thinner when infiltration occurs in Ar.The use of Ar to control the nitride thickness increases the range of possible infiltrants that can be employed, as well as the temperature at which infiltration can occur. A list of the different infiltrants that have been tested is given in Table 1, along with their melting range. Successful infiltration has been possible for all of these infiltrants using a 12h hold at 540°C in N2 to form the skeleton, a 1h purge with Ar and then infiltrating for 2h at 700°C in Ar. The mechanical properties of preforms infiltrated with selected alloys are listed in Table 3. Clearly, the use of lower alloy content infiltrants has resulted in the recovery of some tensile ductility. Nonetheless, the ductility is generally <3%. The infiltrated strength is also largely dependant on the properties of the infiltrant.

Nonetheless, the ductility is generally <3%. The infiltrated strength is also largely dependant on the properties of the infiltrant. Thus, since Mg is the most effective solid solution hardening element for aluminium (14), properties are highest when infiltrating with the Al-5Mg. For infiltrants that are heat treatable, a standard heat treatment cycle provides an increase in strength, but to the detriment of the ductility.Once a ductile infiltrant is used, the embrittling effect of the nitride is more apparent. Figure 9 plots the ductility as a function of nitride thickness (as measured by weight gain) for 6061 – 2%Mg – 1%Sn – 4%nylon preforms nitrided for various times at 540°C. These were subsequently infiltrated with Al – 5%Mg for 2h at 700°C in Ar.

These were subsequently infiltrated with Al – 5%Mg for 2h at 700°C in Ar. Clearly, excessive nitride results in poor tensile ductility and therefore a balance is required between parts containing sufficient nitride to maintain dimensional stability and excessive nitride which degrades the mechanical properties. The addition of Sn and the associated reduction of nitride growth rate makes this balance easier to achieve. When the optimum balance is made between the two, complex parts can be produced with properties equivalent to a common casting alloy, Table 3. Examples of such parts are shown in Figure 10.883.5)%3( y t i l i t c u D e2.521.5l is n e T10.500246810Specimen Weight Gain (%)Figure 9. Effect of amount of nitride (weight gain) on the tensile ductility for 6061 - 2%Mg - 1%Sn - 4%nylon preforms infiltrated with Al - 5%Mg for 2h at 700°C. The embrittling effect of excessive nitride thickness is evident.(a)(b)(c)Figure 10.

The embrittling effect of excessive nitride thickness is evident.(a)(b)(c)Figure 10. Some example parts made by SLS and infiltration, (a) pump body, (b) heat sink, bracket and sculpture and (c) hollow tubes of diameter 35mm (photo courtesy of K. Newall, 3D Systems, Inc).89ConclusionsIn this paper, we have highlighted some of the recent advancements made in the production of infiltrated aluminium components via Selective Laser Sintering. The addition of Sn controls the nitride growth rate. This allows more accurate control of the nitride thickness, which is critical in maintaining ductility and mechanical properties. Tin also improves the infiltration rate, due to the reduction of the surface tension of the liquid as it dissolves the Sn. Finally, switching the process gas to Ar at the end of the nitriding stage not only arrests the nitride growth, but also permits the use of much higher infiltration temperatures, which in turn permits the use of more ductile infiltrants.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAN AEROSPACE INTEGRATED COMPONENT APPLICATION BASED ON SELECTIVE LASER MELTING: DESIGN, FABRICATION AND FE SIMULATIONFei Liua, Wang Jingchaob ,Tan Zhib ,Liu Binb ,Guo Yueb, David Z. Zhanga,*a State Key Laboratory of Mechanical Transmission, Chongqing University, Chongqing 400044, Chinab Beijing Institute of Aerospace Systems Engineering, Beijing 100076, ChinaAbstract: Developments in Additive Manufacturing (AM) technologies have enabled the integrated manufacturing of complex structures with multifunctional performances. Selective Laser Melting (SLM), as one of them, becomes the candidate for production of metallic aviation spaceflight structures because of its high precision for controlling architecture and excellent performance.

An aerospace component named Upper-stage Cabin was selected for integrated design without connectors, and then manufactured by a SLM system using Ti-6Al-4V powder. The dimensional accuracy of the component was verified through testing; It’s mechanical response was analyzed under compressive loading test, compared with the results of numerical simulation. The study found that SLM technology is an effective means for integrated design and manufacture of Upper-stage Cabin structures for aerospace. The dimensional precision of the complex integrated structures formed by SLM meets the application requirements. The formed integrated components have achieved good mechanical properties, which promote the application of this technology in aerospace field.Keywords: additive manufacturing, selective laser melting, upper-stage cabin structure, integrated design and manufacturing1.

IntroductionAerospace components featuring multifunctional performances including complex geometries, lightweight, high-performance and small production runs are difficult, costly, and time-consuming to manufacture. Additive Manufacturing (AM), a layer-by-layer process of joining materials to make objects from 3D model data, is becoming an appropriate solution for these challenges[1,2]. Compared with the conventional subtractive manufacturing methodologies, AM provides unprecedented opportunities for producing complex structures due to its freedom of manufacturing constraints, resulting in simplifying the design and reducing assembly or connection of aerospace components[3]. It meets the requirements of aerospace components with topological complexity resulting from integrated functions.Selective laser melting (SLM), as one of AM technologies for metal forming, has developed rapidly in recent years[4].

Due to its excellent controllability of internal complex structures and good manufacturing accuracy, SLM shows a great capability to topologically488optimize component design with functional integration[5]. It has achieved direct integration in the aerospace field with the application of complex structural components[6]. A substantial weight and cycle reduction can be expected by using SLM for the integrated forming of multifunctional structure[7]. For example, one of the most compelling applications from GE Aviation is the new fuel nozzle fabricated using SLM, achieving the number reduction of parts from 18 to 1, and 25% weight reduction compared to previous component[8].A wide range of materials including titanium and its alloys, nickel-based alloys, stainless and tool steels and some aluminum alloys are available for SLM technology[9].

With the increasing size of parts that can be fabricated by SLM equipment and the continuous expansion of the range of available materials, SLM is showing unique advantages in forming aerospace parts with large-scale and thin-walled complex structures.In space application, the Upper-stage Cabin structure of space vehicle needs to withstand complex load environment, and provide installation and support platforms for the core system's instruments, an example one with complex structures and multifunctional performances is shown in Fig. 1. Therefore it usually exhibits a complex internal structure, at the same time, it is also required to provide good overall stiffness and strength. In addition, there are nearly a hundred machining and sheet metal forming parts installed inside the cabin structure, the manufacturing and assembling of them consume more than 50% of the total cabin cost, as well as the total cycle time.

Integrated design and manufacturing are imperative in order to improve performance, reduce weight, and shorten the manufacturing cycle for aerospace components.Fig. 1 An Upper-stage Cabin in aerospace applicationIn this work, a typical Upper-stage Cabin was selected for integrated design without connectors. In order to study the applicability of SLM technology in aerospace complex structures, a SLM equipment was used to realize the integrated manufacture of scaled Upper- stage Cabin. The dimensional accuracy of the as-build part was verified through testing; Upper- stage Cabin’s mechanical response was analyzed under compressive loading test, compared with the results of numerical simulation.4892. Experimental procedures2.1 Materials and fabricationThe raw material in this work is gas-atomized prealloyed Ti-6Al-4V powders characterized by excellent mechanical properties and high strength-to-weight ratio.

The Ti-6Al- 4V powder have a chemical composition corresponding to ISO 5832-3, as listed in Table 1. Scanning electron microscope (SEM) images of the powders are presented in Fig. 2a and b with different magnifications. These powders are mostly spherical shapes with smooth surfaces, leading to a good flowability and a higher apparent density for SLM process. And they have anarrow particle size distribution with an average size of 33.𝜇𝜇𝜇𝜇Table 1 Chemical composition of Ti-6Al-4V (wt.%) provided by Electro Optical System (EOS) GmbH, GermanyTi-6Al-4VTiAlVONCHFeWt.%(balance) 5.5—6.75 3.5—4.5 ＜0.2 ＜0.05 ＜0.08 ＜0.015 ＜3Fig. 2 SEM micrographs of the Ti-6Al-4V powder, ×100(a) and ×1000(b)The aerospace component was manufactured by a SLM machine (EOSINT-M280, EOS GmbH, Germany) equipped with a 200 W CW Ytterbium fibre laser. The processing wasperformed at an argon protective atmosphere with in the building chamber.

The processing wasperformed at an argon protective atmosphere with in the building chamber. And the processing parameters that have been optimized for the Ti-6Al-4V powder were set up as shown in Table 2. Under these conditions, the density of bulk Ti-6Al-4Vcontent less thanO20.1𝑣𝑣𝑣𝑣𝑣𝑣. %was measured as properties of as-build tensile test piece are shown in Table 3.by the Archimedes method. The mechanical34.42 𝑔𝑔/𝑐𝑐𝜇𝜇(≥ 99.5%)490Table 2 Processing parameters for Ti-6Al-4VParametersValueLaser power, Laser spot size, 𝑊𝑊 Layer thickness, Scanning speed, Hatching space,𝜇𝜇𝜇𝜇 𝜇𝜇𝜇𝜇 𝜇𝜇𝜇𝜇/𝑠𝑠 𝜇𝜇𝜇𝜇175 100 30 125 0.12.2 Test detailsUniaxial compression tests of the Upper-stage Cabin were conducted by using a universal testing machine equipped with a 200KN load cell(RMDW-200 from Ji’nan Rui Machinery Equipment Co., Ltd). The test applies axial pressure on the cabin through a thick steel plate loading tool, as shown in Fig. 6a.

The test applies axial pressure on the cabin through a thick steel plate loading tool, as shown in Fig. 6a. The compressive tests were carried out at a loading speed of 500N/s, accompanied by electronic strain gauges recording the local deformation behaviors of the Cabin.In order to investigate suitability of the simulation method for the production based on the SLM process, the 3D FE analysis with the full sized model were performed for quasi-static simulations using ANSYS FE software. The CAD model of Upper-stage Cabin was imported into ANSYS where the beam element was used to simulate ribs with horizontal and vertical meshes, and the shell element was used to simulate the rest of the skins and the cross-beam structure. Material performance parameters used to this simulation are shown in the Table 3, which were measured by tensile test according to the ISO6892-1:2009 standard.

Material performance parameters used to this simulation are shown in the Table 3, which were measured by tensile test according to the ISO6892-1:2009 standard. The simulation test was performed according to the experimental conditions, and its results were compared with the experimental data.Table 3 Material performance parameters of Ti-6Al-4VPropertiesValueElastic modulus,110Poisson's ratio𝐺𝐺𝐺𝐺𝐺𝐺0.34Density,4.43Yield strength, 𝑔𝑔/𝑐𝑐𝜇𝜇31060Tensile strength,𝑀𝑀𝐺𝐺𝐺𝐺1230𝑀𝑀𝐺𝐺𝐺𝐺3. Results and discussion3.1 Accuracy of Upper-stage Cabin fabricated by SLMThe main load-bearing structure of the cabin was extracted independently, establishing an integrated model with multiple components connected. Specifically, the integrated model is491mainly composed of cross-beam structure and grid-stiffened shells with two ellipse windows, as shown in Fig. 3.

Specifically, the integrated model is491mainly composed of cross-beam structure and grid-stiffened shells with two ellipse windows, as shown in Fig. 3. The cross-beam structure which has been studied in previous literature[10] was applied to realize lightweight design with thin-walled structures. And in order to meet the support-free requirement, chamfering optimization was performed for the horizontal ribs. Due to the size limitations of parts that can be formed in the existing SLM equipment, the model ) of ) of 0.5mm, shell grid’s rib heightwas scaled by 1:12 compared to the actual product, resulting in an overall cabin height ( 120mm, an outer diameter (D) of Ø200mm, a skin thickness ( () of 4mm, and rib width () of 1mm.𝜹𝜹𝑯𝑯𝑹𝑹𝒂𝒂𝑹𝑹Fig. 3 A Upper-stage Cabin structure for aerospace applicationThe finished product is subjected to full-scale geometry measurements. Specifically, the ) of the shell were measured using deep arch micrometer.

Specifically, the ) of the shell were measured using deep arch micrometer. thickness ( The vertical rib width ( ) of the grid, the outer diameter (D) and the height (H) of the shell 𝜹𝜹 are measured using a vernier caliper.

thickness ( The vertical rib width ( ) of the grid, the outer diameter (D) and the height (H) of the shell 𝜹𝜹 are measured using a vernier caliper. Each measurement item was carried out at least 5 times to obtain the average values which are shown in Table 4.) and the rib height (𝑯𝑯𝑹𝑹𝒂𝒂𝑹𝑹Table 4 Comparison of main dimensions of Upper-stage CabinMeasurement itemDesigned size(mm)Measurement size(mean: mm)Error(%)Thickness,0.500.6530Rib height,𝜹𝜹 𝑯𝑯𝑹𝑹Vertical rib width,4.001.003.991.190.2519Outer diameter, D𝒂𝒂𝑹𝑹200.00199.900.5Height, H120.00120.940.78As can be seen from the above table, the studied Upper-stage Cabin produced by the SLM process has a relatively high precision meeting the requirements of aerospace components, indicating the capability of SLM suitble for manufacture of integrated structures with thin- walled features and complex structures.492𝑯𝑯𝑹𝑹Fig.

4 SEM morphologies of the (a)side view and the (b) top view of Upper-stage CabinThe SEM morphologies of the side view and the top view of Upper-stage Cabin areexhibited in Fig. 4. It is found there exist a lot of bonded particles on the side surface of Upper-stage Cabin, while there is almost none on the top surface. For SLM technology, the phenomenon of bonded particles is very common and unavoidable. The main reason for this phenomenon is that part of the powders at the contour boundary are partially melted and adhere to the surface[11]. On the top side of the part, all powders are melted as the laser spot moves, leaving a lot of noticeable scanning traces as shown in Fig. 4b. In our previous and top side work[12], it has been found the roughness of side surface is Ra=40.04 Ra=17.18 particles, scanning traces and layer thickness as well[11]..

4b. In our previous and top side work[12], it has been found the roughness of side surface is Ra=40.04 Ra=17.18 particles, scanning traces and layer thickness as well[11].. The dimensional accuracy errors showed in table 3 come from these bonded𝜇𝜇𝜇𝜇𝜇𝜇𝜇𝜇3.2 Experimental and numerical simulation resultsAs shown in Fig. 5, there were 52 electronic strain gauges arranged on the Upper-stage Cabin, recording the local deformation behaviors of ribs and shell in different positions of cabin.493Fig. 5 a) Arrangement of electronic strain gauges,b) as-build model with electronic strain gauges on one sideFig. 6 a) Compressive experimental test monitored by electronic strain gauges;b) Simulated strain images under 100KN loadingFig. 6a shows the compressive experimental test monitored by electronic strain gauges. In this process, the load was gradually increased from 10KN to 100KN, and the data was extracted every 10KN.

6a shows the compressive experimental test monitored by electronic strain gauges. In this process, the load was gradually increased from 10KN to 100KN, and the data was extracted every 10KN. On the other hand, the numerical simulation method calculated the strain and displacement of the model under various loads. Fig. 6b exhibits the simulated strain images under 100KN loading. As expected, the largest total deformation occurs on both sides of the ellipse window, which were recording by electronic strain gauges multi-directionally.Fig. 7 shows comparison between experimental and numerical results of total deformation at the first loading. It can be seen from the figure that the experimental results and the simulation results are in good agreement overall at both load of 10KN and 100KN, indicating the deformation behavior SLM manufacturing under the quasi-static conditions can be predicted through numerical simulation methods.

Subsequently, Upper-stage Cabin was494subjected to the secondary loading under the stage of elastic deformation. The loading condition is the same as that of the first-time loading. The same results as the first loading are shown in Fig. 8, indicating the feasibility of the simulation method. In addition, like the theoretical design model, the integrated component of SLM manufacturing exhibits good performance consistency at different locations in the structure. This is mainly due to the fact that layer-by- layer processing of additive manufacturing weakens or even eliminates the adverse effect of structural complexity on performance. However, there were some points where the difference between experimental and numerical results is relatively large in both two loadings.

However, there were some points where the difference between experimental and numerical results is relatively large in both two loadings. It was found that most of these points were subjected to large deformations under experimental load, which have gone beyond the scope of linear analysis, so quasi-static linearity no longer applies to this stage.Fig. 7 The comparison between experimental and numerical results of total deformation at the first loading:a) load=10KN; b) load=100KN495Fig. 8 The comparison between experimental and numerical results of total deformation at the secondloading: a) load=10KN; b) load=100KN3.3 Microstrutural analysisThere are many reasons for the error between the experimental results and the finite element simulation results, in which the difference of microstructure inside the parts may be the most critical factor.

In the finite element analysis of the Upper-stage Cabin of this article, the analysis conditions are defined according to the traditional numerical simulation method. In fact, the internal microstructures of parts produced by the SLM are different from the ones produced by the traditional forming method.Fig. 8a and b show the metallographic organization of the traditional rolling part and the as-build SLM manufacturing part, respectively. Fig. 9a,b show the metallographic organization of the traditional rolling part with different magnification. And Fig. 9c,d show the metallographic organization of SLM manufacturing part from the Upper-stage Cabin after test. From figures, it can be seen that metallographic structures of rolling state part exhibit arelatively large lamellar net arrangement characteristic, a typical two-phase structure. The -phase layer. The width of the α-phase . While metallographic structure with elongated grains are shown in Fig.

The -phase layer. The width of the α-phase . While metallographic structure with elongated grains are shown in Fig. 8c and d is about 2 β α for as-build SLM manufacturing part. This is due to the high temperature gradients that occur during the SLM process. The alloy forms fine microstructures in the rapid solidification process,phase flakes are distributed around the coarseα + β𝜇𝜇𝜇𝜇496followed by the martensitic transformation, resulting in a fine acicularstructure.′αFig. 9 Optical microstructure of rolling state(a, b), and as build of SLM(c, d)4. ConclusionIn this study, an integrated component of aerospace Upper-stage Cabin was designed and then manufactured by SLM at a scale of 1:12 using Ti-6Al-4V alloy for aerospace applications. The feasibility of SLM technology for launch vehicle with complex integrated structures was verified in terms of manufacturing precision and mechanical properties.

The feasibility of SLM technology for launch vehicle with complex integrated structures was verified in terms of manufacturing precision and mechanical properties. The experimental and numerical analysis were performed for testing the deformation behaviors under compressive load. It was found that the dimensional accuracy, mechanical properties, and simulation test results are in accordance with engineering practice, providing a basis for the more extensive application of SLM technology in aerospace structural products.

In the future research, with a larger-sized SLM equipment, it will be possible to achieve a real integrated component at a scale of 1:1 designed and formed for aerospace application.The study found that SLM technology is an effective means for integrated design and manufacture of Upper-stage Cabin structures for aerospace, which can significantly shorten the manufacturing cycle and reduce manufacturing processes, thereby reducing manufacturing costs.Although new opportunities and benefits were provided by SLM technologies for manufacturing of aerospace integrated components, there are still some dynamic properties,497such as fatigue life, that need to be further studied.

And with the development of SLM manufacturing in the future work, more accurate and larger size parts to be fabricated will be desirable for aerospace application.Acknowledgments: The authors acknowledge the National High Technology Research and Development Program of China (863 Program: 2015AA042501) for financial support of this work.Conflicts of Interest: The authors declare no conflict of interest.References1.2.Liu, R.; Wang, Z.; Sparks, T.; Liou, F.; Newkirk, J. Aerospace applications of laser additive manufacturing. 2017; p 351-371. Allevi, G.; Cibeca, M.; Fioretti, R.; Marsili, R.; Montanini, R.; Rossi, G. Qualification of additively manufactured aerospace brackets: A comparison between thermoelastic3.4.stress analysis and theoretical results. Measurement 2018, 126, 252-258. Wang, X.; Xu, S.; Zhou, S.; Xu, W.; Leary, M.; Choong, P.; Qian, M.; Brandt, M.; Xie, Y.M.

ile Product Testing with Constrained PrototypesUichung Cho Kristin L. WoodRichard H. CrawfordManufacturing and Design Lab. Department of Mechanical Engineering The University of Texas at Austin Austin, Texas 78712AbstractThe means to acquire reliable functional information is a critical factor that dif ferentiates product development time and cost. Thanks to advances in solid free form industries can produce geometrically complex parts within fabrication techniques, dramatically reduced time and cost. Even though industries can save significant ef forts by performing functional tests rapid prototypes, they still prefer full-scale prod uct tests, especially in later design phases, due to inherent limitations in traditional similarity methods (TSM).

This paper describes a new method to perform reliable functional tests with rapid prototypes that cannot be properly handled by the TSM.1IntroductionThe means to acquire and manage system information, which guides design activities, is influential to the performance of design processes. Our ancestors created primitive devices mainly through natural experiences. Recognizing the need to obtain necessary information with efficiency, people started to perform planned observations (active observations), re flecting on them to exploit specific phenomena. Thanks to such exploits and advances in virtual modeling techniques, contemporary designers are able to obtain necessary functional information far more efficiently.

Thanks to such exploits and advances in virtual modeling techniques, contemporary designers are able to obtain necessary functional information far more efficiently. However, there are still needs to obtain reliable functional information in time, as the phenomena that can be described purely by mathematics are still very limited.Thanks to advances in various rapid prototyping techniques, industries can fabricate geometrically complex parts with dramatically reduced effort (Jacobs, 1992; Aubin, 1994). Considering the significant time and cost spent on product testing to verify or improve final product quality, it is natural to expect various industrial applications of rapid prototypes for function testing.

However, industries utilize rapid prototypes mostly in early design phases, and very limited case studies are reported on functional testing with rapid prototypes to verify product performance or tune design parameters (Dornfield, 1995; O'Reilly, 1993), mainly due to material issues (Wall et al., 1991).

This paper introduces a new method to correlate distorted systems that cannot be handled by traditional methods, especially when systems are composed of multiple parts.292Similarity Methods for PrototypeTo predict full-scale product behavior using economic scaled (rapid) prototypes, mainly two types of similarity methods can be applied:Dimensional analysis that designs scaled models, and correlates product and scaled model behavior on the basis of dimensions of dominant system parameters, through the Buckingham n theorem (Bridggman, 1931; Sedov, 1959; Szucs, 1980; Baker et aI., 1991) .Analytical fractional analysis that mathematically correlates the solutions of two known equations, by comparing and manipulating the equations (Sedov, 1959; Kline, 1966).In dimensional analysis, one should design scaled models so that all of correspondingdimensionless parameters of the scaled model and target system should be identical.

pecially when prototyping materials are limited, the chance to fabricate well-scaled models is very low. Besides, one should precisely know system parameters beforehand. Such lim itations are clarified in our previous publications (Cho et al. 1998(a)(b)). In comparison to dimensional analysis, analytical fractional analysis can provide more accurate scale test ing results (Kline, 1965). However, one should know precise governing equations, material properties, and loading conditions that are not well informed in many circumstances.In order to overcome such limitations in traditional methods, we introduced a novel empirical similarity method (ESM), in which geometrically simple specimens are utilized as additional information sources. One can refer to (Cho and Wood 1997, Cho et al.

One can refer to (Cho and Wood 1997, Cho et al. 1998(a)(b)) for fundamentals of the method, and this paper will describe the concept through an archery bow example focusing on correlation of systems composed of multiple parts.3 System Correlation through Relative State RepresentationWhen products are composed of multiple parts, the TSM requires to design scaled models to satisfy the following material proportionality condition: Proportionality of material properties - The material property ratio of corresponding model and product components should be kept identical, in order to perforrr scale testing with the TSM. The proportionality condition is very difficult to be satisfied, when available prototyping materials are limited. If two systems do not satisfy this proportionality, they are defined30to have distorted configurations, and there exists no method to correlate systems with distorted configurations within the authors' knowledge.

As an initial attempt, we test the possibility to predict behaviors of systems composed of two materials, utilizing a scaled model fabricated from a single prototyping material.The product specimenI , the scaled model, and the product system parameters can be represented in terms of the model specimen2 parameters, through the scale and form factors. One should notice that the form factors and the state scale factors should satisfywhere the state scale factors .AXI and .AX2 can be mathematically derived through dimen sional analysis, and cPm and cPP are unknown form factors. From the equation, the form factors cPm and cPP' which represent the amplifications of the model and product states under pure geometrical changes, should be identical, if .AXI = .AX2' viz.

From the equation, the form factors cPm and cPP' which represent the amplifications of the model and product states under pure geometrical changes, should be identical, if .AXI = .AX2' viz. the model and the model specimen are designed from the same similarity constraints.In our new approach, each part is designed following the IT theorem in order to maintain the identity of the form factors of the model and the product, and the scale factors of the specimen pair and mode-product pair. Through the identical factors, one can represent the lumped model of the product specimen, the scaled model and the product, in terms of the model specimen parameters. As a result, one can represent the product state as a function of the states of the specimens and the model. The step by step procedure of this new approach named as lumped ESM is summarized below.

As a result, one can represent the product state as a function of the states of the specimens and the model. The step by step procedure of this new approach named as lumped ESM is summarized below. Procedure of the Lumped ESM: (1) Build lumped parameter models: Describe the states of the model pair (the model and the model specimen) and the product pair (the product and the product specimen) as lumped models with unknown effective lumped parameters of each part of the systems. (2) Determine the scale factor of each component: Considering the material/loading pa rameters and size scaling, design scaled models through dimensional analysis, and mathe matically derive the scale factors of lumped parameters (the ratio of the product and model parameters) of each component.

(3) Relatively describe lumped models: Employing the derived scale factors and an un known form factor (between the model specimen and the model), represent the product specimen, the scaled model, and the product in terms of lumped parameters of the model specimen. (4) Represent product states in terms of the model and specimen states: By comparing and manipulating the four lumped models, describe the unknown product states as a function of the measured states of the scaled model and the specimen pair.In summary, the lumped ESM correlate the states of model and product pairs by im plicitly deriving the unknown lumped parameters through the measured states of a scaled model, and a specimen pair.lGeometrically simplified version of product 2Geometrically simplified version of scaled model31(1)Figure 1: Exemplary Archery BowsNumerical Archery Bow ExampleIn designing archery bows, the shooting force, viz.

the force required to fully deform a bow to shooting position, is a typical system behavior that draws our attention. One should design bows so that both ergonomic (e.g., required shooting forces) and functional (e.g., arrow speed at a distance) requirements are satisfied.Virtual bow models may not be accurate enough to design quality bows, mainly due to geometrical and material nonlinearity. Large bow deformation causes significant geo metrical nonlinearity, and the stress-strain behavior may not be linear within the wide strain range, considering popular bow materials (e.g., fibers and woods). For these rea sons, designers may need to fabricate several test bows to verify or improve product quality. Considering the geometrical complexity of bow frames shown in Figure 1, significant cost and time to fabricate test-bows are expected. We illustrate and validate the lumped ESM to predict the required (or resultant) force of a product bow in an effective manner.

We illustrate and validate the lumped ESM to predict the required (or resultant) force of a product bow in an effective manner. As a preliminary study for future physical testing, we test lumped ESM using numerically sim ulated bow behaviors. In this bow example, the TSM requires the identity of the Young's modulus ratios of the bow frames and strings of models and products. However, the iden tity of the ratios cannot be satisfied in many cases, as the prototyping material choices to fabricate the string and the frame of model bows are limited.Shooting forces were simulated with ANSYSTM, and the finite element bow models of the target (product and scale model) and specimen bows of interest are shown in Figure 2. In the finite element models, the string is modeled as a two-dimensional elastic link (Spar element), and the bow frame as triangular solid elements (Plane2). The width of the bow frame is set to 35mm, and the cross section area of the string to lOmm2.

The width of the bow frame is set to 35mm, and the cross section area of the string to lOmm2 . As shown in Table 1, we use the same bow string for both model and product bows, considering the difficulty in finding proper frame (or string) materials that satisfy the condition required32FE Models (Grid Size=O.lmXO.lm)(a) Specimen(b) TargetFigure 2: Symmetric Half of the Specimen and Product BowsTable 1: System Parameters of the Bowsby the TSM,where the superscripts band s denote bow frame and string, and the subscripts m and p denote model and product materials, respectively. In this simulation, we consider full-size models for the purpose of clear comparisons.

In this simulation, we consider full-size models for the purpose of clear comparisons. However, one can reduce the size of the model bows without loss of generality, as long as the model systems are designed from the same scale laws.To predict shooting forces of the target fiber bow (product) in Figure 2 (b) through the ESM, it is assumed that bows can be modeled as lumped spring systems shown in Figure 4.

Then, the shooting force of a model specimen (polymer bow with the simple geometry shown in Figure 2 (a)) can be represented as,~ = k~ns' k k b + k smsm,sS8 '33(2)(3)500500g~ u &:400400bO t: 'P 0 0.c "".~ "IX</}0-)002lX))00200100HlJ000.40.600.20.40.6Stting Stretching Distanceo(m)(a) Model Bows(b) Product BowsFigure 3: Simulated Shooting Forces of the Bows+-F,8Figure 4: Lumped Model of a Bow34where is the shooting force (or the force required to deform the bow), 8 is the traveling distance of the bow string, and k b and k S are unknown effective spring constants of the defining a new state variable X, the force equation bow canand string respectively. equivalently represented as8 rmsX ms = Ds = ems + e ,bwhere eb = -\ and eS = represent effective compliance of the bow frame and string.

Then k the states of the scaled model, product specimen, and product bows can be represented asAce~s + <pe~s + eSX psX m X p = Ac<pe~,s + eS,where Aceb ::;- = T is the compliance scale factor, and <p emeb emsform factor of the bow frames.The remaining task is to express X p as X p = h(Xm, X ms , X ps ), so that we can predict X p from the measured states, X m, X ms , and X ps ' From Equations (4) and (5), we can derive the following equations without bow string related parametersX p - X ps = Ac(<p (<p -X m - X ms1)e~s,l)e~,s'By combining the two equations, one can derive the following prediction equation without unknown form factor <p,whereAs a result, one can determine X p (8) from Equation (7) based on the measured X(8) of the specimens and the model, to predict the bow force of the product Fp ( 8) from the definition of the stateFigures 5 (a) and (b) show the shooting force predicted through the TSM and ESM.The TSM cannot predict the bow force with accuracy, as35(4)(5)(6)(7)(8)(9)ROOROO~ 700 ~600II700600tJ..

Of) .~ ..c CIl ] 'S !500400JOO2(8)fIII0II5004<K)JOO2(K)I(K)/10000.20.40.600.20.40.6String Stretching Distance 0 (m)(a) TSM(b) ESM o Force Predicted with the TSM and ESM •Actual Force (ANSYS Simulation)Figure 5: Prediction of the Shooting Force with the TSM and ESMOne way to perform scale testing with the TSM is to assume that the influence of the bow string to the reaction force is relatively small in comparison to the bow frame, In that case, the ratio of the reaction force should be ;;e- = :;-, and it should be equal to 16/7.46 = 2.14. Figure 5(a) shows the force predicted thr01~gh the TSM (Fp = 2.14Fm ), neglecting the effect of the bow strings. As we expected, there exists large discrepancy between the actual and predicted reaction forces of the product bow (Figure 5(a)).

As we expected, there exists large discrepancy between the actual and predicted reaction forces of the product bow (Figure 5(a)). In contrast, the bow force predicted through our new lumped ESM (Equation (7) is plotted in Figure 5(b), and one can notice the remarkably improved prediction accuracy. The ESM results are comparable to the full-scale product testing, and one may save significant effort to fabricate full-scale product bow for testing, especially fabricating scaled models from rapid prototyping processes.4 ConclusionsThrough a simple bow example, the capability of the ESM to correlate systems with dis torted configurations has been demonstrated. In general, specimens are used to estimate specific material properties or determine form factors. In contrast, the ESM explicitly esti mate neither system parameters or form factors.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAGING BEHAVIOR OF POLYAMIDE 12: INTERRELATION BETWEEN BULK CHARACTERISTICS AND PART PROPERTIESK. Wudy*†, D. Drummer*†Institute of Polymer Technology, Friedrich-Alexander University Erlangen-Nürnberg, Am Weichselgarten 6, 91058 Erlangen †Collaborative Research Center 814 – Additive Manufacturing, Friedrich-Alexander University Erlangen-Nürnberg, Am Weichselgarten 6, 91058 ErlangenAbstractThe high process temperatures in combination with long building times during the laser sintering process lead to chemical and physical aging mechanisms on the polymeric feed material. The unmolten partcake material, which acts as a supporting structure, can be removed after each building process and reused for further processes.

The unmolten partcake material, which acts as a supporting structure, can be removed after each building process and reused for further processes. However, material as well as bulk properties are changed due to thermal and mechanical load during the laser sintering process. Within this paper the interrelation between the aging state, bulk values and resulting part properties like porosity, surface roughness and mechanical behavior are derived. Therefore, polyamide 12 powder is used for at least five processing cycles without refreshing. Before and after each building process, bulk characteristics and changes of the particle surface were determined. Specimens were manufactured during the laser sintering process in order to study the part density, roughness and mechanical behavior.IntroductionAdditive manufacturing technologies like selective laser sintering of thermoplastic powders generate components directly from a CAD dataset without needing a form or a mold.

Therefore, the layer wise process of selective laser melting allows the generation of individualized complex parts with a serial character. Contrary to conventional manufacturing techniques like injection molding, which are optimized for high volume production and low unit costs, the costs for additive manufactured parts are almost independent from the degree of complexity and the quantity [1]. The advantages of additive technologies are favored for products with high customization, a high level of complexity and small lot sizes, which is known as “mass customization” [2].Up to now, selective laser melting is widespread in prototype construction and is currently on the threshold of generating serial products in rapid manufacturing applications [1]. For the establishment of such techniques, in rapid manufacturing the main hurdles are a higher process reliability and reproducibility. Both characteristics are directly connected to the material behavior during laser melting process.

Both characteristics are directly connected to the material behavior during laser melting process. Nowadays, merely one third of additive manufactured components are used in engineering applications as functional serial parts [3]. Nonetheless this application area shows the strongest rate of growth during the past years [3]. For this reason, in particular a fundamental understanding of the process itself and process material interaction is indispensable for increasing the amount of serial applications in the sector of additive manufacturing.770State of the artIn selective laser sintering incremental volume units in form of layers are build up in order the generate components without using a mold. Therefore, the thermoplastic polymer powder is selectively fused by a CO2 laser. The process itself consists of the three sub- processes: powder coating, energy input and material consolidation.

Therefore, the thermoplastic polymer powder is selectively fused by a CO2 laser. The process itself consists of the three sub- processes: powder coating, energy input and material consolidation. Firstly, a roller or knife system generates a powder layer in the build chamber, which is tempered to a specific process temperature just below the polymer’s melting temperature. Secondly the cross-section of the component is melted by the use of a CO2 laser, while the surrounding powder particles remain loosely in the build chamber, forming a supporting structure. After exposure, in the phase of material consolidation the build chamber lowers by the thickness of one layer, i.e. 100 µm in most cases and a new powder layer is applied. Step by step, these sub-processes are repeated until the component is completed. Following a cooling phase, the component is then withdrawn from the build chamber.

Step by step, these sub-processes are repeated until the component is completed. Following a cooling phase, the component is then withdrawn from the build chamber. [4, 5]During the laser melting process the thermoplastic polymer powder has to endure a thermal stress, which is determined by the build chamber temperature and the building time. According to the model of quasi-isothermal laser melting process, the build chamber is preheated to a temperature between the melting and crystallization point of the semi crystalline thermoplastic. As a consequence, physical and chemical degradation of the polymer can take place. Physical degradation, a reversible process, changes the order of molecules and leads to post crystallization, concentration changes and agglomeration. Changes in the chemical structure of polymers like chain scission, branching or cross-linking are caused by oxidation, post-condensation and hydrolyses are so called chemical degradation phenomena.

Changes in the chemical structure of polymers like chain scission, branching or cross-linking are caused by oxidation, post-condensation and hydrolyses are so called chemical degradation phenomena. [6-8]Investigations on oven aged polyamide 12 laser sinter powder reveal an increase of viscosity number due to thermally induced post condensation reaction under nitrogen and vacuum atmosphere for short periods of storage time near the beginning of melting [9, 10]. The storage under oxygen for short periods of time leads to an increase and afterwards to a decrease of the average molecular weight due to competing chain branching and scission mechanisms [9, 10]. The influence of storage time and temperature on oven stored polyamide 12 powder on the average molecular weight, the melt volume rate and thermal material properties like melting point was investigated in [11] from Pham.

He shows that the crystallinity decreases and the melt volume rate increases with the storage time and temperature. In addition, Dotchev’s [12] findings confirm the increase of viscosity with increasing time and temperature, probably due to post condensation reaction. Furthermore, Dotchev [12] investigated the correlation between building height, refreshing strategy, powder position and aging state of polyamide powder. Reaction kinetics of the solid state post-condensation of polyamide 12 powder and the prediction of the viscosity dependent on building time and temperature are investigated in [13]. In this paper the authors were able to derive a time and temperature dependent aging model for the process of selective laser melting.Furthermore, previous investigations dealt with the influence of the amount of build processes or rather building time on rheological and bulk material properties as well as on part properties of polyamide 12 [14].

During the first building processes the elongation at break rises significantly, reaches a maximum and drops down sharply afterwards [14]. Gornet [15] also dealt with the influence of the amount of processing cycles on thermal, rheological and mechanical material properties. Wegner et al. [16] compared oven storage experiments with aging during laser melting process in order to establish a simple method to simulate the aging process. Therefore, aging processes at two different sintering systems were performed in order to analyze the correlation between material quality, process parameters and part properties. Josupeit [17] determined the influence of different powder aging states, adjusted with varying refreshing rates on mechanical properties of laser molten parts. Furthermore, Starr et al. [18] analyzed the influence of virgin powder and refreshed or blended powder on mechanical characteristics.

Furthermore, Starr et al. [18] analyzed the influence of virgin powder and refreshed or blended powder on mechanical characteristics. No significant difference between virgin and blended powder can be detected in any mechanical property [18]. Zarringhalam et al. [19] studied the microstructure of laser771sintered parts made of virgin, aged and refreshed powder with microscopic and thermal methods.MethodologyMaterialFor the following investigations an unmodified virgin polyamide 12 (PA12) powder type PA 2200 from the supplier EOS GmbH, Germany is used. In order to obtain significant and reproducible results, new powder material with an equal lot number is used.Processing conditionsThe build processes were carried out at a research sintering system, which was build up by means of the Collaborative Research Center 814 – Additive Manufacturing, which is homed at the Friedrich-Alexander University Erlangen-Nürnberg.

The research sintering system is equipped with a CO2 laser with a wavelength of 10.6 µm, a maximum laser power of 50 watts and a focus diameter of 400 µm from the supplier Synrad. In order to achieve a homogeneous powder surface temperature distribution, the system is provided with a multiple heating system with eight infrared heating zones.Beside the build chamber temperature (172 °C), the energy density (0.35 J/mm³), the laser power (7.8 W), the scan speed (904 mm/s), the layer thickness (100 µm), the powder application speed (250 mm/sec), the hatch distance (250 µm) as well as the building height (75 mm) were held constant during the building process.

A constant number of six tensile test specimens were produced within each manufacturing process to avoid changes of the thermal load.The correlation between thermal stress time or rather number of processing cycle and aging behavior or the partcake powder, as well as on resulting part properties like mechanical behavior, density and surface roughness were determined. Therefore, virgin PA12 powder was processed five times without refreshing. Approximately only 60 per cent by weight of the feed material of a building process becomes partcake material. The other 40 per cent is used for component generation or is transferred to the over flow container. By means of this at the beginning ten equal manufacturing processes were built with virgin PA12 powder at the beginning. The sieved and mixed (30 minutes, 400 revolutions per minute) powder from these processes form the basis for further processing cycles.

The sieved and mixed (30 minutes, 400 revolutions per minute) powder from these processes form the basis for further processing cycles. The accumulated thermal stress time the powder has to endure can be calculated from the period of each building process. Additionally, the temperature of the powder bed surface is observed by a pyrometer during the whole manufacturing process. The detailed design of experiments is shown in table 1.Table 1. Design of Experimentsprocess cycle1 2 3 4 5build temperature (°C) 172 172 172 172 172build time (h) 8.4 8.4 8.4 8.4 8.4cumulative build time (h) 8.4 16.8 25.2 33.6 42.0772Characterization of material propertiesAfter each build process material properties, e. g. bulk density, Hausner ratio and viscosity number were analyzed in order to determine the process relevant aging mechanism. According to DIN EN ISO 60 [20] the bulk density after and before each build process is measured.

According to DIN EN ISO 60 [20] the bulk density after and before each build process is measured. The bulk density is defined as the mass of the polymer particles divided by the total volume they occupy. For the analysis a bulk density tester from Emmeran Karg Industrietechnik type ADP is used. The average was calculated out of at least five measurements.The Hausner ratio, a dimensionless parameter to characterize the compressibility of bulk materials, is defined as the quotient of tap and bulk density. The tap density was analyzed according to DIN EN ISO 787-11. Contrary to the norm the densification was realized via manual tapping. For this work a standardized number of at least 100 taps over a period of 5 min was used. The experiments were carried out under ambient conditions. Five runs build the basis for calculation of average Hausner ratio and root mean square deviation. According to Abdullah [21] the Hauser ratio indicates the flowability of powder materials.

Five runs build the basis for calculation of average Hausner ratio and root mean square deviation. According to Abdullah [21] the Hauser ratio indicates the flowability of powder materials. The best flowability is obtained when a bulk material does not condense. In this case the tap density equals the bulk density and the Hausner ratio is one. With increasing Hausner ratio the flowability of the bulk material declines.In the laser melting process the bonding of two layers and thus the mechanical properties are fundamentally dependent on the viscosity of the polymer melt. A high melt viscosity leads to insufficient bonding, which results in low mechanical properties. To characterize the changes in the viscosity and thus the average molecular weight due to solid state post-condensation, the viscosity number of the virgin and aged powder is determined. The viscosity number of PA 12 is usually determined with m-Cresol as a solvent.

The viscosity number of PA 12 is usually determined with m-Cresol as a solvent. Due to security and health issues sulfuric acid at 25 °C is used in the following investigations.Characterization of part propertiesIn order to set up a correlation between powder properties and resulting mechanical characteristics intermediate values like porosity and surface roughness have to be determined. The component density was analyzed by determining the specimen dimensions and weight. The dimensions of the specimen for density measurements are 10 x 10 x 4 mm. Out of five specimens the average density was calculated.The surface roughness of the tensile test bars was measured with a laser scanning microscope type LEXT OLS4100 form Olympus Corporation. A measurement field of 650 x 650 µm is resulting with a lens with 20x magnification.

A measurement field of 650 x 650 µm is resulting with a lens with 20x magnification. The area related average surface roughness SZ utilized at three different positions at the final manufactured layer of the test specimen.According to DIN EN ISO 527-1, the test specimens’ mechanical properties are determined in tensile tests at a test speed of 5 mm/min. The test speed for the determination of the Young's modulus is 1 mm/min. The 5 test specimens are examined to their elastic moduli, maximum stresses and elongations at break. The samples are dried at 70 °C in a vacuum oven in order to create constant boundary conditions.

The 5 test specimens are examined to their elastic moduli, maximum stresses and elongations at break. The samples are dried at 70 °C in a vacuum oven in order to create constant boundary conditions. The moisture content of the samples after the oven storage is measured with Karl Fischer titration and reached a value of 0.2 wt.-%.The morphology and inner structure of the manufactured components were illustratedon 10 µm cross sections via transmitted light microscopy.773Results and discussionBulk properties like, bulk density and Hauser ratio of the basic material affect the mechanical behavior of laser molten parts. In addition, bulk properties are dependent on the aging state of the used polymer powder. Hence, bulk density and Hausner ratio are determined for powder that has endured varied thermal stress time in selective laser sintering process. Figure 1 represents the bulk density and the Hausner ratio for increasing cumulative build time.

Figure 1 represents the bulk density and the Hausner ratio for increasing cumulative build time. The bulk density decreases with rising thermal stress time, whereby the Hausner ratio shows a vice versa trend. A consequence of the lower bulk density will be a reduced specimen density. The rise of the Hausner ratio goes along with reduced powder flowability and thus a poor surface quality of the coated layer. Both mechanisms might influence the mechanical properties of the laser sintered parts.Figure 1. Bulk density and Hausner ratio of aged PA12 powder dependent on cumulative build timeThe reason for the reduced bulk density are manifold. There might be a change in the particle size or in the particles surface for example. Therefore, Figure 2 shows the volumetric particle size distribution of the aged powders.Figure 2. Particles size distribution of aged powder dependent on cumulative build timeWith increasing cumulative build time, the particles size distribution is unchanged.

Particles size distribution of aged powder dependent on cumulative build timeWith increasing cumulative build time, the particles size distribution is unchanged. Hence, the theory that changed particle size is responsible for the reduced bulk density can be excluded. Mielicki [22] confirms these results. For powder oven aged at a temperature of 174 °C neither 774the particle size nor the particle form changes with increasing storage time [22]. Nevertheless, for high storage times the deviation of form factors like the aspect ratio rises, which indicates variation in the particle form or surface [20]. Beside unchanged particle sizes the curves show a narrow particles size distribution, which is favored for high bulk and thus high component densities.Beside shifted particles size distribution, changes of the particles surface could be responsible for the decrease of bulk density with increasing cumulative build time.

Commercially available PA12 powder like Duraform PA from 3d Systems is adhered with fumed silica to enhance particle flowability and thus generate higher bulk densities [1]. The adhered fumed silica on the particle surface acts as spacer and reduces the adhesion force between the particles. Therefore, the flowability and the bulk density rises. Changes of the concentration of fumed silica may affect the powder flowability. That is why the particle surface is shown in Figure 3 via scanning electron microscopy images with a magnification of 1,000 and 10,000.magnification1,000 x10,000 xr e d w o p n i g r i ve mr e d w o p d e g ai t d l i u b e v i t a l u m u c () h 0 . 2 4Figure 3. Scanning electron microscopy pictures of virgin and aged PA12 powder particlesWith a magnification of 1,000 between the virgin and aged powder no differences can be obtained. A closer look at the particles surface reveals significant differences of the two powders.

A closer look at the particles surface reveals significant differences of the two powders. For high thermal stress times in a laser sintering system the particle surface is smooth and a fewer amount of adhered nano scaled particles, assumed as fumed silica, is visible. Both effects, the changes in particles surface and the reduction of fumed silica evoke that the distance between the particles is less and thus the adhesion force will be enlarged. Higher adhesion forces signify lower bulk densities and a reduced flowability. Hence, the increase of the bulk density can be traced back to a change of particle surface. However, the reduction of powder flowability will presumably affect the process stability and reproducibility in a negative way and the increasing bulk material porosity will lead to less dense parts.775Figure 4.

Viscosity number of aged PA12 powder dependent on cumulative build timeBeside reduced bulk density rheological characteristics like viscosity number changes with the thermal stress time, Figure 4. The materials’ viscosity is fundamental for the bonding between two layers and hence for the mechanical properties of the laser molten parts. With increasing cumulative build time, the viscosity number rises until a steady state is reached. The rise of the viscosity number can be in all likelihood traced back to post condensation processes shown for example in [14]. The average molecular weight rises due to post condensation reaction, where stored monomers react with chain ends with a high molecular mobility. According to the Mark-Houwink-Equation also known as Kuhn-Mark-Houwink-Sakurada- Equation (Equation 1) with increasing molecular weight the intrinsic viscosity strives against a limit [23].

According to the Mark-Houwink-Equation also known as Kuhn-Mark-Houwink-Sakurada- Equation (Equation 1) with increasing molecular weight the intrinsic viscosity strives against a limit [23]. This correlation is valid for a structural coefficient α between 1 and 0.6, which is the case for polymer solutions [23].Equation 1𝛼𝛼[𝜂𝜂] = 𝐾𝐾 ∙ 𝑀𝑀With increasing viscosity, the sinter behavior as well as the bubble removal time in selective laser sintering is changed. According to the Frenkel’s theory of sintering with increasing viscosity the time for particle coalescence rises [24]. Additionally, the bubble removal time in a polymer melt increases for rising viscosity, which is known from rotational molding [25].

Additionally, the bubble removal time in a polymer melt increases for rising viscosity, which is known from rotational molding [25]. Both effects affect the time until a homogeneous melted layer in laser sintering is formed and thus influences the component density.Therefore, the component density is evaluated for different cumulative build times of the used basic powder in order to find a correlation between bulk and part properties dependent on the aging time, Figure 5. With increasing thermal stress time of the PA12 powder the density of the produced parts decreases. At the one hand the reduction of part density is related with the lower bulk density. On the other hand, due to higher viscosity and molecular weight the bubble removal time and the coalescence time rises, which indicates a higher bubble concentration at the end of a building process.

On the other hand, due to higher viscosity and molecular weight the bubble removal time and the coalescence time rises, which indicates a higher bubble concentration at the end of a building process. Investigation on polyethylene particles via rotational molding reveals that for decreasing melt volume rates and increasing average molecular weight, the time for the disappearing of trapped air will be longer [26]. Both effects, the lower bulk density and the higher viscosity for longer thermal stress times of the PA12 basic powder signify a reduced part density, which is represented in Figure 5.776Figure 5. Component density and surface roughness of laser molten specimen built with aged PA12 powder with varying cumulative build timeIn addition to component density the average surface roughness Sz of the via laser sintering produced specimen is influenced by the cumulative build time of the used basic powder, Figure 5.

For high cumulative build times of the used powder the surfaces roughness rises from 80 µm to approximately 120 µm, which is an increase of 50 %. With increasing viscosity of the melt the resistance to embed particles on the top layer is heightened, which may cause different surface roughness.Both the reduced component density and the raised surface roughness may influence mechanical properties. Pores as well as embedded particles act as defects and stress peaks, which triggers fracture. In the case of rapid manufacturing applications, resulting mechanical part properties of the laser molten parts are of main importance. Hence, Figure 6 represents the tensile tests of specimen produced with powder with different thermal stress time.Figure 6.

Hence, Figure 6 represents the tensile tests of specimen produced with powder with different thermal stress time.Figure 6. Representative tensile test curves and elongation at break of laser molten specimens build with aged PA12 powder with varying cumulative build timeFigure 6 shows representative stress strain curves of laser sintered specimens produced with PA12 powder, which has endured different thermal stress times. With increasing cumulative build time, the stress strain curves show a significant change in fracture behavior. An initially ductile material behavior changes with increasing powder aging to a brittle performance. This transformation in fracture behavior is also visible by analyzing the elongation at break dependent on cumulative build time of the used PA12 powder. The elongation at break increases up to a powder aging state of the feed material of 16 hours and 777drops down sharply afterwards.

The elongation at break increases up to a powder aging state of the feed material of 16 hours and 777drops down sharply afterwards. Figure 4 reveals that with increasing cumulative build time the viscosity number und thus the average molecular weight increases, probably due to post- condensation reaction. This increase of molecular weight due to storage at temperatures near the beginning of melting of the polymer leads to longer distances in which the chains can slide past each other and thus the elongation at break raises. For cumulative build times higher than 20 h the bulk density decreases sharply. This goes along with increasing part porosity and surface roughness. As a result of all these effects the elongation at break drop down. Finally, the increasing porosity and surface roughness are superimposing effects and mechanical part characteristics collapse.

As a result of all these effects the elongation at break drop down. Finally, the increasing porosity and surface roughness are superimposing effects and mechanical part characteristics collapse. Furthermore, microstructural changes, that are not analyzed yet, may occur and have an influence on the mechanical behavior.Figure 7. Elastic modulus and max. tensile strength of laser molten specimen build with aged PA12 powder with varying cumulative build timeFigure 7 exhibits the mechanical properties elastic modulus and maximum tensile strength as a function of the powder aging state of the feed material. Both characteristics, the elastic modulus and the tensile strength, are not influenced by the powder aging state of the used basic material.

Both characteristics, the elastic modulus and the tensile strength, are not influenced by the powder aging state of the used basic material. The abidance on a stable level is not surprising because both characteristic values do not act as a feasible indicator for the aging state of the material [6].Figure 8 shows transmitted light microscopy images of laser sintered polyamide 12 specimen. The cumulative build time of the basic powder for specimen production rises from the left to right side. Specimen build with virgin powder represent pores and unmolten particles, which exhibit under polarized light in a yellowish appearance. Furthermore, the build direction is visible due to layer wise appearance. As a result of high number of unmolten particles within the component the elongation of break reaches a lower level than for the specimen, which has endured a higher thermal stress. Nevertheless, the unmolten particles have almost the same density as the raw material.

Nevertheless, the unmolten particles have almost the same density as the raw material. That is why the density of the parts is almost constant during the first build processes. Notwithstanding, the unmolten particles act as defects and elongation at break is reduced.With increasing cumulative build time of the basic powder the unmolten particles disappear and the single layers are no longer visible. A rising amount of molten material signifies an increase in the elongation at break. Further increase of the cumulative build time of the basic powder leads to a high amount of huge pores, a reduction of density and thus reduced mechanical characteristics.778Figure 8. Transmitted light microscopy images of laser molten specimen build with aged PA12 powder with varying cumulative build timeConclusionsWithin this paper the influence of build time or rather cumulative build time on the aging state of polyamide 12 paper and its influence on mechanical properties are investigated.

Therefore, material is reused until it has been going through five processing cycles without refreshing. Before and after each laser melting process, the bulk and material properties are analyzed. In addition, mechanical part characteristics are determined in order to correlate them with the powder aging state.With increasing cumulative build time, the bulk density decreases and the Hauser ratio increases by constant particle size distribution. This change of bulk properties with increasing aging time can be traced back to changes of the particles surface due to partial melting or incorporation of the nano scaled additive on the particle surface. In addition, viscosity increases due to high temperature during laser melting process with increasing cumulative build time.

In addition, viscosity increases due to high temperature during laser melting process with increasing cumulative build time. Probably this effect is caused by post-condensation reaction of the polyamide.Furthermore, a decline of the part density and an increase of the surface roughness can be detected with increasing cumulative build time. These superimposing effects lead to a change in the elongation of break. With increasing aging time of the used feed powder, the elongation at break passes a maximum and declines sharply afterwards. Due to the increase of the viscosity number and the reduction of the amount of unmolten particles the elongation at break rises initially. The reduction of the bearing cross section due to higher porosity, in combination with higher surface roughness, leads to the decrease of elongation at break.

-POWERED AUGMENTED REALITY TRAINING FOR METAL ADDITIVE MANUFACTURINGDonald Palomino1,2, Shijie Bian1, Pedro Salcedo2, Erick Navarro3, Bharath Gopalaiah2, Richard Otis4, Bingbing Li1,2 1Autonomy Research Center for STEAHM, California State University, Northridge, CA 91324 2Department of Manufacturing Systems Engineering and Management, California State University, Northridge, CA 91330-8332 3SimInsights Inc, Irvine, CA 92618 4Jet Propulsion Laboratory, California Institute of Technology, Pasadena, CA 91109, USAAbstract Metal additive manufacturing (AM) training can cost a company the considerable amount of time, cost, and resources. To resolve this challenge, the NSF funded HyperSkill platform was applied to create an innovative, immersive training program that integrates AI-powered object detection and text recognition into a comprehensive digital twin of the metal AM machine operation.

This immersive digital twin will support the delivery of just-in-time guidance to trainees while also monitoring their actions and providing contextual and personalized feedback to accelerate training, foster retention, and maximize transfer to the actual job. The augmented reality (AR) training supports the import of 3D assets, no-code authoring of workflows, standard operating procedures (SOP), step-by-step instructions, and delivery across a wide variety of AR devices. For this study, the specific metal AM operation is based on the Renishaw AM400 in both its full and reduced build volume (RBV) configuration.Keywords: Augmented Reality Training; Metal Additive Manufacturing; Text Recognition; Object Detection1.Introduction Metal AM is a developing technology with huge potential in major industries like theaerospace, defense, and healthcare sectors.

While relatively limited in build volume relative to traditional manufacturing, metal AM is capable of producing parts with intricate geometries and detail unobtainable by CNC machining and die casting. Additionally, unlike subtractive manufacturing, metal AM eliminates the need for expensive tooling and excessive waste. Metal AM is also able to print from a variety of different metals, including titanium which is notoriously known to be difficult to machine conventionally. From heat exchangers with complex patterns to biomedical micro-structures, metal AM is flexible and robust enough to fulfill the needs of any manufacturing facility. Due to these benefits, metal 3D printers have seen a sharp increase in market share in recent years with the global metal 3D printing market expected to grow from $378.0 million in 2019 to $738.8 million by 20251.As the number of metal 3D printers proliferate, so too does the number of employeesneeded to operate these printers.

Currently, the training process requires an individual experienced in the operation of the metal 3D printer to teach the trainee. This experienced individual is either a veteran operator at the company, or a professional trainer from the metal 3D printer’s company requested by the trainee’s company to come on-site. In both cases, this is a major cost as either the company will have to temporarily sacrifice the work productivity of the965veteran operator, or they will have to pay thousands of dollars for the on-site professional trainer. In addition to this, the training process can take weeks and requires the use of a metal 3D printer for demonstration and training purposes, leaving the printer out of commission for its intended use. These costs from time, money, and resources quickly add up to become a significant expense for companies in adopting metal AM.

These costs from time, money, and resources quickly add up to become a significant expense for companies in adopting metal AM. Lastly, this training process makes it inconvenient for companies to own multiple metal 3D printers from different manufacturers, as a veteran operator specializing in one printer would be unable to teach a trainee how to operate another printer, making companies resort to having to pay for an on-site professional trainer.A system that is capable of standardizing the metal AM workflow in a comprehensiveand easy to understand manner would greatly alleviate these associated costs for companies. The system would also give companies more flexibility in the different types and brands of metal 3D printers they can use.In this study, an Artificial Intelligent (AI)-powered AR training simulation for metal AM was developed to simplify the training process and reduce its costs.

The NSF funded Hyperskill platform will be applied to create an innovative, immersive training environment that will integrate AI-powered text recognition and object detection into a comprehensive digital twin of the metal AM processes. This immersive digital twin will support the delivery of just-in-time guidance to trainees while also monitoring their actions and providing contextual and personalized feedback to accelerate training, foster retention, and maximize transfer to the actual job.

The AR training supports the import of 3D assets, no-code authoring of workflows, standard operating procedures, step-by-step instructions, and delivery across a wide variety of AR and VR devices.For this case study, the digital twin will be created using Solidworks and Fusion 360CAD software, the Renishaw AM400 metal 3D printer in both its full and RBV configuration will be used as the digital twin’s real-world counterpart, and the HoloLens 2 will be used as the AR headset to immerse the trainee in the training environment. An RBV is a printer add-on that allows it to produce smaller builds relative to its full configuration. The smaller build volume takes up significantly less resources than the latter, requires less metal powder to print, and reduces the amount of time it takes to change between different metal powders.1.1. Current Status of Metal AM, AR Training, and AI Systems1.1.1.

Current Status of Metal AM, AR Training, and AI Systems1.1.1. Metal AMDespite the major advantages that metal AM offers, it currently suffers from a number ofissues. One of these problems is the lack of a workforce with the proficiency and qualifications necessary to operate metal 3D printers2. This stems from the larger issue of a deficiency in the general education for AM. Universities and training programs have not kept up with the development of new AM technologies. Therefore, more training, education, and awareness needs to be accomplished for the current workforce and the new generation of students3. Progress towards this is already underway. For example, the Center for Innovative Materials Processing through Direct Digital Deposition at Penn State offers combined lecture and hands-on training in their AM facilities. However, they only trained 240 industry practitioners over three years due to966space constraints and safety concerns4.

However, they only trained 240 industry practitioners over three years due to966space constraints and safety concerns4. Additionally, several universities and companies including Virginia Polytechnic Institute, Virginia Tech, University of Texas Austin, Society of Manufacturing Engineers, and Stratasys are implementing courses and training specializing in AM5. However, despite these efforts, it is not fast enough to meet the growing demand of metal AM. They also only cover AM in a broad scope, not the specific operation of metal 3D printers. The need for a cost-effective and versatile training program for metal AM still exists.1.1.2. AR TrainingAR technology has proven itself to be effective in providing more capabilities to workers and in training. Through case studies performed by Microsoft on their HoloLens2, the same AR headset used in this study, it was found that a 177% return on investment can be achieved by implementing the AR headset into their manufacturing process6.

AR technology has been used in the automotive industry to assist drivers in replacing a flat tire by tracking the car wheel and relevant tools7. Based on the work done in implementing AR technology to a training simulation for an industrial setting, it was found that tasks asked of the trainees must be of a high enough complexity to make the AR system feel justified. Additionally, “selling” the benefits of the AR system to the trainees in its improved work efficiency and build-in error control is important in the trainee’s reception to it8. An overall advantage in using AR technology in training is the integration of multiple digital formats and capabilities into one function process. Even complex assemblies can be broken down into easy to follow steps that an inexperienced employee can follow9.1.1.3. AI SystemsAI systems are capable of recognizing both body and hand gestures of a human. Thesegestures can be used as a means of communication between the human and computer10.

AI SystemsAI systems are capable of recognizing both body and hand gestures of a human. Thesegestures can be used as a means of communication between the human and computer10. This can eliminate the need for peripheral devices like a mouse and keyboard which limit a user to the desk on which they are placed. Implementing AI systems into AR or VR simulations allows the user to work in a virtual environment while being able to freely stand and walk around, not being constrained by fixed input devices. AI-powered object detection can be used in developing a real-time system to monitor machine states and perform automatic fault detection on 3D printers11. Using character-region awareness for text detection (CRAFT)12, AI systems can also identify and recognize the text and finger positions of a user operating a 3D printer with a control panel. This is done by processing images recorded on a wearable camera13.9672. AI Algorithm2.1. Text Recognition2.1.1.

This is done by processing images recorded on a wearable camera13.9672. AI Algorithm2.1. Text Recognition2.1.1. Contour DetectionFigure 1: Flowchart for contour recognitionBefore the substantial step of text recognition, preprocessing procedures are required tolocalize the text regions to narrow down the search space and discard the noisy background contexts. Considering that a majority of text displayed on the screen panel are bounded within a colored region, a contour detection step is performed to segregate them. As shown in Figure 1, the input image is first transformed into its grayscale representation after being re-scaled to a suitable resolution, followed by a binary pixel threshold of RGB value (0, 255, 150) and a neighborhood-oriented adaptive threshold screening to find contours.

Utilizing OpenCV’s functionalities of calculating contour area and perimeter, the contours that map to a closed region are filtered out and considered the regions of interest for the subsequent text detection step.2.1.2. Text Region LocalizationCharacter region awareness for text detection (CRAFT), a machine learning model thatsupports the localization of characters and phrases, is adopted to restrict the search space for text recognition further. Specifically, the regions of interest obtained from the previous step are passed through a pre-trained CRAFT model, which localizes individual characters within the said region by evaluating the probability that a given pixel is the center of the character. Subsequently, CRAFT merges localized character regions into instances by inference of the probability of adjacent characters belonging to the same word.

Subsequently, CRAFT merges localized character regions into instances by inference of the probability of adjacent characters belonging to the same word. The output of CRAFT consists of the bounding box coordinates of localized text regions, which are then passed to the final text recognition module. Note that the contour-detecting preprocessing step that narrows down the regions of interest, together with CRAFT which empowers accurate detection of text regions, allows the proposed framework to function effectively in diverse working environments.9682.1.3.

Recognition AlgorithmThe localized text regions of interest, as represented by bounding box coordinates, arepassed through a unified four-stage scene text recognition framework, which includes a thin- plate spline (TPS) for feature transformation, a pre-trained residual network (ResNet) for feature extraction, a bi-directional long short-term memory (BLSTM) for sequence modeling, and an attention mechanism for prediction generation. The output from the recognition framework consists of the predicted texts for each region of interest, along with the corresponding prediction confidence score. The predicted texts are used in subsequent steps for the real-time identification of the machine’s operation phase and the instructions as displayed on the output screen.2.2.

The predicted texts are used in subsequent steps for the real-time identification of the machine’s operation phase and the instructions as displayed on the output screen.2.2. Object DetectionOne traditional approach for detecting the operator’s interaction with the machine would be to use automated finger detection models to locate the position of the operator’s fingers on the interactive panels. However, directly adopting and integrating a finger detection model into the proposed framework will likely render the framework cumbersome, which may affect the real- time performance during work-scenario testing. Therefore, a color-change method is devised to utilize the localized text regions of interest together with the detected closed contours as generated from the previous text recognition section. Specifically, for each detected closed contour region, the averaged pixel color is calculated and compared with a pre-defined metric.

Specifically, for each detected closed contour region, the averaged pixel color is calculated and compared with a pre-defined metric. If a finger intrudes the region, the algorithm would be able to automatically detect the anomaly in region colors, thereby identifying the location of the finger.9693. Creation of AR Training3.1. Creation of Digital Twin using CAD SoftwareFigure 2: a) Renishaw AM400 with accompanying hardware, b) Build area in full configuration, c) Build area in RBV configuration3.1.1.Full ConfigurationThe first step in developing the AR training simulation is to use computer aided design(CAD) modeling to create a digital twin of the metal 3D printer. All components of the metal 3D printer relevant and necessary to its manual operation were created, as well as components for facilitating ease in part recognition.

All components of the metal 3D printer relevant and necessary to its manual operation were created, as well as components for facilitating ease in part recognition. Parts hidden behind panels such as electronics and other hardware components used in machine maintenance were not created.To begin the CAD design process using Solidworks, components that are critical to theoperation of the metal 3D printer had their measurements taken manually with calipers and970measuring tape. This was done to represent the digital twin as close as possible to its real-world counterpart. This includes the metal 3D printer’s build area and its internal components such as the build plate, the wiper, and the wiper adjustment knobs. Unique to the full configuration is the build plate which is a large, flat, and solid piece of metal.

Unique to the full configuration is the build plate which is a large, flat, and solid piece of metal. Some of the non-critical parts had their measurements approximated as they are not required for the general operation of the metal 3D printer and only serve as visual representations to aid in the learning experience. An example of this was the computer as not every feature it included was necessary for operation, only the on/off switch. The safe-change filter area is another example of this since although many components are visible within the area, most are not needed in general operation, such as the pumps, generators, and compressors. The metal 3D printer also requires the following accompanying hardware in its operation process: the chiller, the dryer, the immersion separation vacuum, and the argon tank. As such, these components were also made into CAD models.

As such, these components were also made into CAD models. The full digital twin model of the Renishaw AM400 and its accompanying equipment is shown in Figure 2a), and a close up of its build area in full configuration is shown in Figure 2b).A key part of the AR training simulation is highlighting parts relevant to the current step. To accomplish this, all interactable components like the knobs, handles, and buttons were created as separate parts using Solidworks, then added to their respective CAD models as an assembly. Once the CAD assembly was finished, the model was converted into a file that was compatible with HyperSkill. This was done by using an add-in called Extended Reality (XR) Exporter in Solidworks. This converted the default .SLDPRT files into the .glb (extended reality) files required by HyperSkill. These .glb files were then imported into HyperSkill to complete the digital twin of the Renishaw AM400 in its full configuration.3.1.2.

These .glb files were then imported into HyperSkill to complete the digital twin of the Renishaw AM400 in its full configuration.3.1.2. RBV ConfigurationTo create the digital twin of the metal 3D printer in its RBV configuration, a similarprocess to its full configuration is performed. All components of the metal 3D printer relevant and necessary to its manual operation were created, as well as components for facilitating ease in part recognition. Parts hidden behind panels such as electronics and other hardware components used in machine maintenance were not created.Unlike the full configuration digital twin process, Fusion 360 is the software used tocreate the CAD models for the RBV configuration. Similar to the full configuration process, all components that are critical to the operation of the metal 3D printer had their measurements taken manually with calipers and measuring tape.

Similar to the full configuration process, all components that are critical to the operation of the metal 3D printer had their measurements taken manually with calipers and measuring tape. This includes the metal 3D printer’s build area and its internal components such as the RBV unit, the wiper, and the wiper adjustment knobs. Unlike the full configuration’s build plate, which was large, flat, and solid piece of metal, the RBV unit is much more complex and consists of several components including the RBV recoater, cover, adjustment blocks, top plate, cover plate, build plate, and powder compacting tool. Similar to the previous process, some of the non-critical parts had their measurements approximated as they are not required for the general operation of the metal 3D printer and only serve as visual representations to aid in the learning experience.

The following accompanying hardware to the metal 3D printer were also made as CAD models using Fusion 360: the chiller, the dryer, the immersion separation vacuum, and the argon tank. The full digital twin model of the RBV configuration is identical to the full configuration with the exception of the build area,971which features an RBV build plate as opposed to a full build plate. A close up of its build area in RBV configuration is shown in Figure 2c).Similar to the full configuration process, all interactable components were created as separate parts using Fusion 360, then added to their respective CAD models as an assembly. Once the CAD assembly was finished, an add-in for Fusion 360 was used to convert the default .F3D files into the .glb files required by Hyperskill. These .glb files were then imported into Hyperskill to complete the digital twin of the Renishaw AM400 in its RBV configuration.3.2.

These .glb files were then imported into Hyperskill to complete the digital twin of the Renishaw AM400 in its RBV configuration.3.2. Training Procedure and Flowchart3.2.1.Full ConfigurationFigure 3: High-level flowchart for full configurationFigure 4: Low-level flowchart for full configuration detailing loading the full build plateIt is critical to the training simulation’s effectiveness that the training procedure for operating the metal 3D printer in its full configuration is as comprehensive and easy to follow as possible. To create the training procedure, it is necessary to be trained by an experienced operator. During this training, each step is documented in detail and video of the process is captured for reference. From these notes, two step-by-step procedures can be created: one high- level and one low-level.

From these notes, two step-by-step procedures can be created: one high- level and one low-level. The high-level procedure lists 10 generalized steps, as shown in Figure 3, while the low-level procedure breaks each of these steps into many more and in far greater detail. A segment of the low-level procedure is shown in Figure 4. This will then be reviewed by the experienced operator to ensure that all of the steps were documented correctly and that the steps were in the proper order. Once the training procedure has been verified, it can be implemented into HyperSkill.9723.2.2. RBV ConfigurationFigure 5: High-level flowchart for RBV configurationFigure 6: Low-level flowchart for RBV configuration detailing turning on all the hardwareSimilar to the creation of the training procedure for the metal 3D printer’s fullconfiguration, it is also critical that the RBV configuration procedure is comprehensive and easy to follow.

As an experienced operator goes through the operation, each step is documented in detail and video of the process is captured for reference. Two step-by-step procedures are then created based on these notes: one high-level and one low-level. While both the full and RBV configuration procedures are similar to each other, they differ greatly in loading the build plate and metal powder. The high-level procedure for the RBV configuration is shown in Figure 5, while a segment of its low-level procedure is shown in Figure 6. This will then be reviewed by the experienced operator to ensure that all of the steps were documented correctly and that the steps were in the proper order. Once the training procedure has been verified, it can be implemented into HyperSkill.9733.3.

Once the training procedure has been verified, it can be implemented into HyperSkill.9733.3. HyperSkill ProcedureFigure 7: a) Hierarchy tree within HyperSkill, b) Assembly tab, c) Simulation with digital twinHyperSkill, developed by SimInsights, is a no-code 3D simulation software for both ARand VR purposes. HyperSkill enables users with little to no programming experience to create immersive training content. It also enables non-programmers to author AR and VR content and to collect and visualize experience data. It features a cloud service to store interactive virtual objects converted from existing CAD and 3D assets.After the converted .glb files of the metal 3D printer’s digital twin are imported intoHyperSkill, they are then opened within the Virtual Objects tab of the software. To the left side of the panel is the hierarchy tree, as shown in Figure 7a).

To the left side of the panel is the hierarchy tree, as shown in Figure 7a). When selecting a feature in the hierarchy tree, its corresponding object or surface in the digital twin will be selected and turned bold. The features within the hierarchy tree can also be renamed for convenience when assigning them to steps in the training procedure.On the right side of the panel is the assembly tab, as shown in Figure 7b). The assemblytab is where the step-by-step instructions for the training simulation are written. Available options for defining the assembly are mechanical, wiring, and snapshot. For this study, only the mechanical feature will be used. Each step added to the assembly tab is numbered and kept in numerical order. When creating a new step, first the title and instructions must be added. There is an option to add media, such as images or videos, that can be viewed during the training simulation to aid the trainee.

There is an option to add media, such as images or videos, that can be viewed during the training simulation to aid the trainee. Lastly, features within the hierarchy tree that are used in the current step are selected. These selected features correspond to its respective object in the digital twin and will highlight it during the step within the training simulation. This helps to direct the trainee’s attention to areas of significance regarding their current step. Consider adding an example as a figure.Once all the steps necessary to the operation of the metal 3D printer have been completed, the Virtual Objects tab can be exited and the Simulations tab is selected. A new simulation is created by giving it a name and choosing a premade scenario as the virtual environment. This name is important as it will be used to later call the training simulation within the AR headset.

This name is important as it will be used to later call the training simulation within the AR headset. Lastly, the virtual object with the .glb file and step-by-step instructions are974simply dropped into the simulation as shown in Figure 7c). The training simulation is now ready to be used within the AR headset.3.4.Implementation of AR HeadsetFigure 8: Digital twin of Renishaw AM400 being scaled down to fit user environmentAfter the digital twin is created and imported into the HyperSkill simulation, it is ready tobe imported into the AR headset. First, the HyperSkill app must be installed. Because the software is considered 3rd party and not affiliated with Microsoft, the manufacturer of HoloLens 2, HyperSkill must be manually downloaded. This is done by first connecting the AR headset to a computer via the AR headset’s personal IP address. The device then needs to be put into developer mode, found within the onboard settings of the HoloLens 2.

The device then needs to be put into developer mode, found within the onboard settings of the HoloLens 2. Being in developer mode allows the user to add 3rd party software and apps to the device. The HyperSkill app is then downloaded from the connected computer to the AR headset. Once it has been downloaded, the computer may be disconnected and the HoloLens 2 can then be worn. Within the menu of the AR headset, the HyperSkill app can be opened, which then prompts the user to select the simulation they would like to run. After selecting a simulation, the training commences with the digital twin appearing in a 1:1 ratio. Depending on the size of the user’s environment, they may wish to scale the digital twin down to a more convenient size, as can be seen in Figure 8.

Depending on the size of the user’s environment, they may wish to scale the digital twin down to a more convenient size, as can be seen in Figure 8. Once an appropriate scale is set and its positioned anchored, the training simulation will begin.9753.5.Integration of AI-Powered Text Recognition and Object DetectionCapture image from cameraSend image to edge devicePre-process imageGenerate AI predictionsSend predictions to Hololens deviceProcess predictionFigure 9: High-level flowchart of prediction processThe AI system will interface with the HyperSkill HoloLens app through arepresentational state transfer application programming interface (REST API). The REST API will be deployed on a local edge device using a Python Flask web application. We use a Flask application instead of deploying the AI system on the Hololens due to the limited battery and processing power of the Hololens.

We use a Flask application instead of deploying the AI system on the Hololens due to the limited battery and processing power of the Hololens. Deploying the AI system on an edge device further simplifies the process of validating the AI performance and accuracy. Should there be a need, the Flask application can also be deployed in the cloud at scale.One the AI system is deployed on the edge device, the HoloLens application will be configured to communicate with the flask web app. After a certain set of frames or at the request of the Hololens user, HyperSkill will capture an image from the front-facing Hololens camera. The image will be sent to the flask web app through the REST API. The web app will then process the image and feed it into the AI system. The AI system will then process and image and return the corresponding predictions. The predictions are sent back to the Hololens device for further processing.

The AI system will then process and image and return the corresponding predictions. The predictions are sent back to the Hololens device for further processing. A high-level flowchart of the prediction process of the AI system once integrated into the HyperSkill HoloLens app is shown in Figure 9.9764. Discussion4.1. Accuracy of AI-Powered Text Recognition4.1.1. Baseline: all lights on scenarioFigure 10: Baseline condition with all lights on; a) Raw image frame, b) Image frame with predicted phase and textThe accuracy of the text recognition for the AI system is an extremely important metric in creating a smooth, immersive experience for the trainee. It allows for the training simulation to recognize the screen and phase that the trainee is on using text recognition. If poor accuracy by the AI system occurs, this may result in frustration over the AR training simulation which would be detrimental to its effectiveness.

If poor accuracy by the AI system occurs, this may result in frustration over the AR training simulation which would be detrimental to its effectiveness. Therefore, it is vital to the project that text recognition is as accurate as possible.To test the accuracy of the text recognition, a video of all available phases for the Renishaw AM400 was taken. To ensure the screen panel, as well as the reset button and emergency off button were in frame at all times, a tripod was used for stabilization. A user pressed buttons on the screen to go through each phase used during normal operation of the printer. The user also paused for a few seconds in between each button press to allow an unobstructed view of the screen for the AI system to later process. To set a baseline for future testing scenarios, all lights in the room were turned on and the operator did not wear any gloves during filming. Once all screens and menus were captured on video, it was separated by individual frames.

Once all screens and menus were captured on video, it was separated by individual frames. These images were then manually processed to remove images that were blurry, showed a blank screen during phase transition, or had the user’s hand blocking the screen. Images with the user’s hand in the frame were relocated to a separate folder for future testing of object detection. The remaining images were then processed by the AI system to recognize the current phase and all text on the screen, as shown in Figure 10. Successful recognition is identified when a red box is formed around text. Once all unobstructed frames of the video have been processed, manual verification by a user was then performed to compare the results by the AI system to the ground truth.977Once the accuracy of the AI system’s ability to recognize text was verified to be acceptable under nominal conditions, it is then tested under different working conditions and scenarios.

These external factors in the environment can affect the AI system’s performance during the operation of the AR simulation. Such factors include low-light conditions in the training room or area and harsh light reflecting off the metal 3D printer screen.4.1.2. Low-light ScenarioFigure 11: Condition with half-lights on; a) Raw image file, b) Image frame with predicted phase and textLow-light conditions may affect the ability for the AI system to recognize text. Whilemost factory environments are required to have a sufficient amount of footcandles determined by their local law, shadows from equipment or structures may be cast on the area of the metal 3D printer. Additionally, in more rare instances, the metal 3D printer may be located in a room with below-regulation lighting conditions. In both cases, low-light conditions are simulated to gauge the effect it has on the AI system.To test low-light conditions, a video of all available phases for the Renishaw AM400 wastaken.

In both cases, low-light conditions are simulated to gauge the effect it has on the AI system.To test low-light conditions, a video of all available phases for the Renishaw AM400 wastaken. Similar to the baseline test, a tripod is used to capture the screen, reset button, and emergency off button. However, unlike the baseline test which had all lights in the room on, this test had only half of the lights on. Once all screens and menus were captured on video, it was separated by individual frames based on the screen shown. These images were then manually processed to remove images that were blurry, showed a blank screen during phase transition, or had the user’s hand blocking the screen. Images with the user’s hand in the frame were relocated to a separate folder for future testing of object detection. The remaining images were then processed by the AI system to recognize the current phase and all text on the screen, as shown in Figure 11.

The remaining images were then processed by the AI system to recognize the current phase and all text on the screen, as shown in Figure 11. Successful recognition is identified when a red box is formed around text. Once all unobstructed frames of the video have been processed, manual verification by a user was then performed to compare the results by the AI system to the ground truth.9784.1.3.Screen Reflection ScenarioFigure 12: Condition with harsh, direct light; a) Raw image file, b) Image frame with predicted phase and textThe screen of the metal 3D printer is highly susceptible to reflection as it has no anti-glare film or treatment. While the reflection of the user does not interfere with the manual operation of the printer, it may be an issue for the AI system’s ability to detect the trainee’s hand or recognize text. The screen of the metal 3D printer does not come with an anti-glare film, making it susceptible to reflection.

The screen of the metal 3D printer does not come with an anti-glare film, making it susceptible to reflection. Both excessive lighting and the operator can be reflected off the screen, adding noise and interference to the AI system’s image processing.To test the effect of harsh reflection off of the screen, a video of all available phases forthe Renishaw AM400 was taken. All lights in the room were kept on, as well as two bright flashlights pointed directly towards the screen. While the video recorded, the flashlights would slowly move around the screen, making sure to go over the text. Once all screens and menus were captured on video, it was separated by individual frames based on the screen shown. These images were then manually processed to remove images that were blurry, showed a blank screen during phase transition, or had the user’s hand blocking the screen. Images with the user’s hand in the frame were relocated to a separate folder for future testing of object detection.

Images with the user’s hand in the frame were relocated to a separate folder for future testing of object detection. The remaining images were then processed by the AI system to recognize the current phase and all text on the screen, as shown in Figure 12. Successful recognition is identified when a red box is formed around text. Once all unobstructed frames of the video have been processed, manual verification by a user was then performed to compare the results by the AI system to the ground truth.9794.1.4.Scenario Test ResultsTable 1: Text recognition accuracy results for test scenariosTest ScenarioPhase Identification AccuracyText Recognition AccuracyText Recognition ConfidenceBaseline (All lights on) 100%94.24%0.7180% (± 0.086%)Half-lights on88.83%89.94%0.7098% (± 0.091%)Reflection91.43%91.37%0.6923% (± 0.069%)Recorded in Table 1 are the test results for the text recognition module.

Note that all the numerical results are recorded in the format of average ± standard deviation. Specifically, three test scenarios are adopted to simulate the diverse working conditions of the machine: the baseline scenario in which the camera is fixed and all lights in the room are on; a low-light scenario in which the camera is fixed and only half the lights in the rooms are on; and a screen reflection scenario where harsh, direct light flow moves around the screen causing glare. For each scenario, numerous image frames are extracted from the captured test videos and then run through the text recognition framework to generate text predictions. The predicted texts for each image frame are utilized to identify the operation phase that the machine is in, and the final phase identification accuracy for each scenario is calculated as the proportion of correctly predicted phases over the total number of image frames in that scenario.

For each frame of each scenario, the proportion of correctly predicted texts of interest over the total number of texts of interest is calculated to be the text recognition accuracy of that specific frame, which is averaged together with the recognition accuracy of all other frames to be the final accuracy for that scenario. The confidence score is obtained in a similar procedure.The test results show that the baseline scenario's averaged performances for phase andtext recognition are the most satisfactory. Specifically, when abundant light is provided, the image as captured by the camera shows the most apparent contour distinctions, serving as a premise for subsequent detection and recognition steps. Since most real-life working scenarios support normal lighting conditions, the test results assure the effectiveness of the proposed design. Performances under insufficient or excessive illumination are less satisfactory.

Performances under insufficient or excessive illumination are less satisfactory. When the light intensity is low, text that is close together and bound inside dark-colored display regions tend to be particularly challenging to correctly identify due to the difficulty for the CRAFT model to localize the regions of interest. Likewise, when harsh lights are on, significant reflections on the display screen will affect the text region localization step as performed by CRAFT, thereby reducing the accuracy and confidence of text and phase recognition. Therefore, these working scenarios should be avoided if possible. The confidence scores follow a similar pattern. In the baseline scenario, where the camera's movement is minimal and the lights are abundant, the text prediction confidence is the highest. In comparison, the scenarios in which only half of the lights are provided reflect less confidence in the text detection results.

In comparison, the scenarios in which only half of the lights are provided reflect less confidence in the text detection results. Harsh lighting conditions that result in strong reflections also restrict the text detection performances, which is something to avoid or ameliorate.9804.2. Validation of IntegrationThe complete AR training simulation comprises both hardware and software including:the digital twin, the AR headset, and the Hyperskill app. It is essential that all elements integrate in a cohesive way for the best user experience. Before integration is performed, each system was individually tested to ensure that they were performing as intended. Once each system is verified, a full run of the training simulation was performed.To verify that the digital twin is functional, all interactable components that are referenced in the training procedure, like the knobs, handles, and buttons, are checked within the digital twin assembly.

Additionally, all non-interactable components, like the body of the metal 3D printer, are checked to make sure that it visually resembles the real-world printer to aid in the trainee’s experience. This was done by opening the hierarchy tree within HyperSkill’s Virtual Object tab, then going through the individual parts with the training procedure to ensure all used components have been made and are included.To verify that the AR headset is operational, it first must be able to turn on and recognize basic gestures. These include gestures to open the main menu and to select buttons close up and far away. If the AR headset is unable to recognize any one of these gestures, then it will not be capable of running the AR training simulation. As an additional measure, the calibration app will also be used to adjust the AR headset to the user’s eyes.

As an additional measure, the calibration app will also be used to adjust the AR headset to the user’s eyes. These are basic features of the AR headset and should be highly reliable in its standard operation.To verify that the HyperSkill app is working properly, the digital twin is first checked tosee that it is successfully able to be dropped into the HyperSkill’s virtual simulation on a computer. Once this is achieved, the HyperSkill app on the AR headset is checked for its ability to open correctly. Lastly, the training simulation is selected and its ability to progress through its step-by-step instructions while highlighting relevant components is checked. This procedure requires the verification of the previous systems and results in the final integration of one system for the AR training simulation. After the simulation has been run through in its entirety with no errors or bugs, testing with volunteers can be conducted to gauge reaction and learning effectiveness.4.3.

After the simulation has been run through in its entirety with no errors or bugs, testing with volunteers can be conducted to gauge reaction and learning effectiveness.4.3. Validation of AR Training EffectivenessAfter the training simulation and its individual components are validated, its effectiveness must also be validated. To accomplish this, the Kirkpatrick evaluation model was used. Normally this is a four step process, however as the last two steps pertain to evaluating results over a long period of time, only the first two steps will be utilized for this project. This first step is Reaction: to what degree do participants react favorably to the learning event. The second step is Learning: to what degree do participants acquire the intended knowledge, skills, attitude, confidence, and commitment based on their participation in the training14. For this study, two volunteers participated in the training simulation.

For this study, two volunteers participated in the training simulation. These volunteers were chosen based on having an engineering background with no prior experience in operating a metal 3D printer or using an AR headset.981Table 2: Results of Reaction Survey After Completion of AR Training SimulationQuestions (scored 1-5; 1=strongly disagree, 5=strongly agree) Volunteer1Volunteer 2Training simulation was enjoyable45AR technology is exciting55AR finger and hand gesturing is intuitive44Step-by-step instructions were easy to follow55Teaching style is time-effective45For this first step, Reaction, a proctor administered the training simulation to the participant and assisted when necessary. Prior to beginning the simulation, a tutorial by the AR headset is taken to familiarize the volunteer with how to interact in the augmented environment and with finger and hand gestures for navigation.

An eye calibration was also performed to ensure the headset was optimized for each volunteer. Afterwards, the AR training simulation was given to the volunteers. Upon completion of the simulation, which took around 30 minutes, a survey was conducted to gauge the degree to which the participant found the training favorable and engaging. The results of the survey for both volunteers are shown in Table 2. As can be seen, both volunteers found the training simulation to be enjoyable, intuitive, easy to follow, and time-effective. Verbal feedback was positive and mostly praised the novelty of AR technology and its potential.For the second step, Learning, a short quiz was given to the volunteers before and afterthe training simulation to gauge their retention of the information. All of the questions were simple in nature, mainly locating points of interest within the metal 3D printer.

All of the questions were simple in nature, mainly locating points of interest within the metal 3D printer. However, the initial test scores were low, with one volunteer scoring a 0/5 and the other scoring a 3/5. This shows the complexity of the printer, especially for someone inexperienced in its operation. After the AR training simulation was completed, the volunteers took the same quiz again, however this time the volunteers scored a 4/5 and a 5/5 respectively. Therefore, the simulation indeed improved the volunteers' knowledge of the printer’s operation.5. Conclusion and Future WorkThe rate at which companies invest in metal AM far exceeds the existing education andawareness for a capable workforce to operate metal 3D printers. Conventional methods in hiring new operators involve intensive, long, and expensive training, costing these companies noticeable expenditures.

Conventional methods in hiring new operators involve intensive, long, and expensive training, costing these companies noticeable expenditures. To solve this problem, an AI-powered AR training simulation for metal AM was developed to simplify the training process, reduce its associated costs, and reduce its time frame. This was achieved by using CAD software to 3D model a digital twin of a metal 3D printer, importing it into HyperSkill with the printer’s training procedure, then running the training simulation using an AR headset. Leveraging advanced machine learning models, an end- to-end workflow for the automated processing of images mimicking the operator's vision is developed. This workflow utilizes contour detection and character-region awareness for the982localization of regions of interest, enhancing the subsequent modules' robustness in removing the unnecessary background details within captured images.

Subsequently, a unified four-stage scene text recognition module is adapted to perform text recognition on the previously filtered regions of interest. Utilizing the detected texts, the AI system integrated into the application is capable of recognizing the machine's current phase and the text displayed on the machine's screen panel, all done in real-time and automation, without any human-labored processing. Test results obtained from numerous scenario settings reflect the competent performance of the proposed methodology to function in diverse environments. Feedback using the Kirkpatrick evaluation model showed that volunteer reaction to the AR training simulation was positive and excited at the prospect of AR technology.

Feedback using the Kirkpatrick evaluation model showed that volunteer reaction to the AR training simulation was positive and excited at the prospect of AR technology. It also showed that the volunteers' knowledge test scores improved significantly, a testament to the learning efficiency of the training simulation.Future work may be done to further enhance the robustness and accuracy of text detectionby integrating rectification through similarity search with the aid of the predicted screen phase. Specifically, a data set of mappings between the phases to the ground-truths texts displayed during that phase should be constructed. After obtaining the text predictions through real-time inference, the results are compared to the possible ground-truth texts recorded in the dataset as mapped from the detected phase, and the most similar ground truths are retrieved, which may serve as the rectified results of the predicted texts.

Another aspect of future work would be combining the screen phase and text detection with object and finger detection, which empowers the algorithm to fully supervise and analyze human-machine interactions. If realized, the AI system for real-time automated inference can be integrated with existing AR technologies for vision capturing, thereby producing AI-powered systems to serve as an auxiliary unit to support the operators’ operations of complex devices. Additionally, these capabilities will be expanded to not only include screen phase and text recognition, but also recognize user interaction with the following areas of the printer: filter area, build area door, glove door, laser chiller, dryer, argon tank, and vacuum cleaner. Once implemented, the AR training will be capable of identifying and verifying the actions of a user with any metal 3D printer.6. AcknowledgementsThis research was supported by the U.S.

Once implemented, the AR training will be capable of identifying and verifying the actions of a user with any metal 3D printer.6. AcknowledgementsThis research was supported by the U.S. NASA MUREP Institutional ResearchOpportunity (MIRO) program (80NSSC19M0200), NASA MUREP High Volume (80NSSC22M0132), and partially carried out at the Jet Propulsion Laboratory, California Institute of Technology, under a contract with the National Aeronautics and Space Administration (80NM0018D0004).7. References[1] “Metal 3D Printing Market Accounted for $378.0 Million in 2019 and Is Expected to Be $738.8 Million by 2025: Global Metal 3D Printing Market Anticipated to Reach $738. 8 Million by 2025. Market Report Coverage - Metal 3D Printing. Market Segmentation.” NASDAQ OMX’s News Release Distribution Channel, NASDAQ OMX Corporate Solutions, Inc, 2020.[2] Vafadar, Ana, et al.

OJ Ceramics Made by CAM-LEM (Computer-Aided Manufacturing Engineering Materials) TechnologyJ.D. Cawley, P. Wei, Z. E. Liu, W. S. Newman, B. B. Mathewson*, and A.HeuerCase School ofEngineering Case Western Reserve University 10900 Euclid Avenue Cleveland OH 44106-7204CAM-LEM Inc. Glenville Enterprise,. Center 540 E. 105th St. Cleveland, Ohio 44108AbstractThe concept ofCAM-LEM technology is presented and discussed in the context of thefabrication of Ah03 ceramics. Particular attention is paid to the interplay of green tape characteristics and the unit operations involved in CAM-LEM. Examples of ceramic shapes difficult to form by conventional methods are described.IntroductionCeramic materials offer a wide range of attractive engineering properties.

Examples of ceramic shapes difficult to form by conventional methods are described.IntroductionCeramic materials offer a wide range of attractive engineering properties. For example:their thermal conductivity can be very low (e.g., vitreous Si02) or very high (e.g., AIN); some are excellent electrical insulators (e.g., AI203), others are semiconductors (e.g., SiC) or show metallic conductivity (e.g., TiB2). Structurally, many show good corrosion resistance, and, increasingly, ceramics are being developed with high fracture toughness and strength (e.g., partially-stabilized Zr02 and ShN4). In particular, the properties of advanced ceramics have been markedly improved in the last decade as the result of intensive research into the relationships between powder synthesis, processing and properties [1].Nevertheless, shape forming remains a critical issue that is always challenging.

Conventional ceramic forming methods are either of limited precision or require a prohibitive investment in tooling, or both. Diamond machining can produce many shapes, but is expensive [1] and generally compromises mechanical properties [1,2]. It is also true that many desirable shapes cannot be fabricated using conventional methods.CAM-LEM technology is a novel Solid Freeform Fabrication (SFF) method developed to allow direct fabrication of components of nearly arbitrary complexity using engineering ceramics of arbitrary composition without the need for tooling. The roots of the CAM-LEM approach lie the fabrication of multi-layer ceramic substrates for microelectronic packaging [3]. In suchsubstrates, complex internal wiring is created by stacking thin sheets of ceramic powder distributed in a porous polymer matrix (so called "green tape") that have been punched and screen printed with powdered metal inks.

The processing of such substrates, while representing some of the most advanced ceramic processing, is geared toward mass production of a singledesign. CAM-LEM is a true SFF technique capable of producing arbitrary one-of-a-kind components. (While attention is restricted to ceramics in this paper, we are currently extending the technology to the production of both metallic and engineering plastic components.)The CAM-LEM Process for Engineering CeramicsThe feedstock for the CAM-LEM process included two types of sheet material: green ceramic tape and a fugitive tape. Green ceramic tape is widely available, as it is the basis of a multi-billion dollar microelectronic packaging business. As green tape, the properties are dominated by the polymer matrix; the ceramic powder acts, for the most part, as an inert filler. This allows a wide variety of ceramics to be formed using one of several tape forming processes (tape casting, roll compaction, etc.).

This allows a wide variety of ceramics to be formed using one of several tape forming processes (tape casting, roll compaction, etc.). Fugitive tape is manufactured to exhibit viscoelastic properties that are similar to those of the green ceramic tape, but is completely burned out of the 3D structure during firing.The CAM-LEM system generates a green ceramic component that must be postprocessed and sintered to form a dense monolith. The details of the machine design will be presented elsewhere [4]. One distinguishing feature ofCAM-LEM technology is that a "cut then stack" approach is used. This allows sections cut from multiple materials to be used within a given layer of the stack, if desired. It also allows pieces of complex internal geometry to be fabricated, as shown in Fig 1.Figure 1. Three layer part fabricated using CAM-LEM technology; a) is a room lit view and b) is a back lit view.

Three layer part fabricated using CAM-LEM technology; a) is a room lit view and b) is a back lit view. The piece is designed as a flow channel with an inlet in the top left corner of the upper surface (visible in a) connected by a complex flow channel in the central sheet (visible in b) to an outlet in the bottom left corner of lower sheet.The central steps in the process associated with fabrication of a complex part are schematically illustrated in Fig. 2. There are two critical components in the apparatus (which is discussed elsewhere in these Proceedings [4], a cutting station and a stacking platform.

2. There are two critical components in the apparatus (which is discussed elsewhere in these Proceedings [4], a cutting station and a stacking platform. An arbitrary component is built using the following steps: 1) a sheet of green ceramic tape is picked up robotically and placed on the cutting station; 2) the desired outline is cut using a CO2 laser; 3) waste is selectively removed and the cut piece is transferred robotically onto the stacking table; 4) a sheet of fugitive tape is similarly processed to yield a complementary piece (either inside, outside, or both) and the fugitive is robotically added to the stack and its waste removed.

The cycle is repeated until the 3-D object is finished, complete with the fugitive supports.Green Cer\l.mlc TapeThe finished part is 5) laminated to fuse the individual layers, 6) heated to bum out both the fugitive and the binder in the green tape and finally, 7) the component is sintered to complete densification.~4""'...a!B~ Green Assembly PlatformAssembleI Cut next layer ~Stack on previous layerFugitive Organic TapeTop ViewAs Stacked! Fire to removefugitive layers and sinter green tape.The critical issues for fabricating ceramic components using CAM-LEM technology are: i) suitability of the tape for laser cutting; ii) development of an efficient method for "tacking" during assembly (necessary to preserve registration during handling); iii) post-assembly lamination to create a monolithic green part; and iv) preservation of dimensions during sintering.

Our observations are consistent with recent reports describing the effect of the organic phase on the physical and mechanical properties of green tape [5], and that flaws in and between tapes can be healed during thermocompressive lamination [6].Repeat process until object is completeFinished partFigure 2. Schematic outline of the CAM-LEM process building a green ceramic component surrounded by, and cored with, a fugitive tape that burns out during fIring.Experimental ResultsMaterialsTwo types of green ceramic tape were employed in our experiments, a commercial 94% Al20 3 tape (Coors Electronic Materials, Chattanooga TN) and an in-house formulation based on AI6-SG alumina (Alcoa, Pittsburgh PA). The Coors tape was cast from a nonaqueous slurry with polyvinyl butyrate as the primary binder.

The Coors tape was cast from a nonaqueous slurry with polyvinyl butyrate as the primary binder. The in-house tape was cast from an aqueous slurry with a combination of acrylic emulsion binders (Rohm and Haas, Springhouse, PA); in some cases, polypropylene was added as a plasticizer. The compositions of several in-house tapes are given in Table 1.The principal difference between the two types of green ceramic tape was their respectiveglass transition temperatures. These were determined by a dynamic mechanical analyzer to be 40°C for the Coors tape and -20°C for the in-house tape.11In addition, a novel fugitive tape was developed based on the use of corn starch as a filler. Chemical grade corn starch (Sigma Chemicals, St. Louis, Mo.) has a reported loss on ignition (or ash content) of<O,4%; thus it can be expected to burn cleanly from the component during firing.

Chemical grade corn starch (Sigma Chemicals, St. Louis, Mo.) has a reported loss on ignition (or ash content) of<O,4%; thus it can be expected to burn cleanly from the component during firing. Aqueous-based slurries were initially employed, but the dried tape adhered strongly to the mylar, presumably because the starch is slightly soluble in water and served as a glue during drying. Nonaqueous slurries using toluene and polyvinyl butyrate released easily from the mylar. Samples of the three tapes are shown in Fig. 3.Figure 3.

"'Dispersant (Rohm and Haas) .........Acrylic Emulsion Binders (Rohm and Haas) ............Polypropylene GlycolLaser CuttingThe effectiveness oflaser cutting is a complex function of the nature of the surface supporting the green tape and the composition of the green tape (both the type of polymer and the composition of the ceramic).12The Coors tape was not designed for laser cutting and proved a challenge. Two problemswere encountered. Firstly, detritus that recondensed after the laser was employed for cutting caused undesirable adhesion of the cutout piece to the cutting station platen. Secondly, when cutting thick tape (e.g., nominally 24 mil, or 600 ).lm), a significant taper was observed. Both problems were reduced in severity through a redesign of the cutting station platen, [4] and components could then be automatically fabricated using this tape.The in-house tape proved to be better suited for Jaser cutting.

Adherence to the platendoes not occur and edge taper is much less severe. For these reasons, development of the in house formulation continues. (This tape also shows promise in its ease of lamination.) However, the fabrication experiments reported here concentrate on results obtained with the Coors tape.TackingThe glass transition temperature of the Coors tape is 40°C and the recommendedlamination temperature for this tape is 80°C. The use of thermal tacking during build would require complex thermal management and such high temperatures might significantly soften the entire assembly, to the point where gross flow is likely. Instead a solvent-based approach was employed.The principle of solvent lamination is to locally and temporarily dissolve the polymer onthe surfaces to be mated, so that polymer chains in adjacent tapes become entwined.

With porous tapes, the solvent is redistributed by capillary action, but in all cases it is ultimately evaporated and the polymer returns to the solid state.Often the solvent used is the same as that employed in the original slurry. In ourexperiments, however, ethanol (rather than toluene) was most effective. Ethanol wets and spreads rapidly on the Coors tape, so that application using a dropper was entirely adequate.The solvent was applied to the top surface of the existing stack prior to the placement of each cutout. Timing and quantity of solvent proved important, as excessive solvent can migrate into the cutout, causing adherence to the robot, or insufficient solvent will not be present to soften the surface of the cutout. With correct timing, tall stacks (e.g., 40 layers) were easily produced. This tacking procedure produced robust pieces, and several components were directly fired after tacking and yielded sound ceramics.

This tacking procedure produced robust pieces, and several components were directly fired after tacking and yielded sound ceramics. Nevertheless, for strength critical parts, it is desirable to post-process to assure complete lamination efficiency.LaminationIsostatic pressing is widely employed in ceramic forming. Typically, the part is placed inan impermeable bag, which in then evacuated, sealed, and sublnerged in a hydraulic pressure vessel.

Typically, the part is placed inan impermeable bag, which in then evacuated, sealed, and sublnerged in a hydraulic pressure vessel. Often the fluid is at room temperature (so called cold isostatic pressing), but it can be heated (warm isostatic pressing).13Experiments using a commercial warm isostatic press (National Forge Co., Andover, Ma.)to fabricate bend specimens confirmed that lamination efficiency is easily achieved.Isostatic pressing can be very cumbersome if the piece is delicate or of complex shape.One important role for the fugitive tape is to assist in pressure transmittal to complex components during isostatic pressing; absent the fugitive component, isostatic pressing might lead to deformation of delicate portions of complex structures.FiringThe two issues arising during firing are the ease of binder removal and the extent to whichshrinkage can be controlled. Standard firing schedules were used to successfully remove the binder.

Standard firing schedules were used to successfully remove the binder. No special procedures were required.Shrinkage issues are comparable to those encountered using powder compacts derivedfrom conventional ceramic forming operations. For example, a part of relatively complex geometry was cut from a single green tape; after firing, the relative dimensions were preserved to 1%, see Fig 4.abcFigure 4. Digital image of a) a green part cut from a single sheet of green tape and b) the fired piece. The firing shrinkage was nominally 17%. unage c is a difference image obtain by subtracting the green image scaled to 83% from the image of the fired piece. (Black and white shading reveal differences in the two images.) The degree of registration is overall very good, less than 1% variation being independently measured for the width as a function of position along the length.

The small anisotropy (1%) in shrinkage is also visible~ note the white fringes lie outside left-to-right, but are inside up-and-down.Test ComponentAs a test of the CAM-LEM method, the software and machine design, as well as the compatibility of CAM-LEM technology to ceramics, a virtual part was rendered in the 94% Al20 3 tape. The test piece was a CAD file taken from the AutoCAD (AutoDESK, Sausalito CA) tutorial. The CAD file was translated to STL and computationally sliced. Sheets of Coors tape were then laser-cut and robotically stacked using ethanol tacking, yielding the green compact14shown in Fig. 5a. This green compact was fired using the following schedule: 5 hour ramp to 600°C, 4 hour at 600°, 3 hour ramp to 1560°C, 0.5 hours hold at 1560°C, and power off cool. The fired piece is shown in Fig. 5b.Figure 5. Component derived from an AutoCAD file. Image a) was taken of the green assembly after tack lamination and image b) was taken after firing.

The fired piece is shown in Fig. 5b.Figure 5. Component derived from an AutoCAD file. Image a) was taken of the green assembly after tack lamination and image b) was taken after firing. The edge length of the base on the fired piece is nominally two inches.SummaryCAM-LEM, a novel SFF technology optimized for the direct fabrication of componentsfrom engineered ceramics, has been conceptualized, developed, and demonstrated.AcknowledgementsThis work was supported by the Army Research Office (STTR Contract #DAAH04-94-C-0055), the Office ofNaval Research (Contract #N00014-95-1-0107), and the National Science Foundation (Grant #DMI-94-20373). In addition, resources were also made available through the Cleveland Advanced Manufacturing Program (CAMP).15References1.B. Aronson, "Can We Get Serious About Ceramics," Man. Eng., p. 54, August (1993).2.4. D. W. Richerson, Modern Ceramic Engineering, Second Edition, Marcel-Dekker, New York, (1992).3.

2O3 Precursor Evaluation for SALD Joining C. M. Weiss, H.L. Marcus Institute of Materials Science, Materials Science and Engineering Program/CMBE, University of Connecticut, StorrsAbstract:SFF technology is not limited to the creation of components and objects. Freeformtechnology can be a mechanism for joining of materials. By defining the space between two objects as the boundary for a free-form object it is possible to form a three dimensional joint fill. Selective Area Laser Deposition or SALD, has been used for free form and joining of carbides and nitrides but has the potential to make other classes of ceramics. By selecting a metal-organic precursor and an oxidizer, an oxide can be created by CVD gas decomposition.

By selecting a metal-organic precursor and an oxidizer, an oxide can be created by CVD gas decomposition. In this study such oxide joint fill precursors are studied, particularly for alumina.Background:Gas phase laser chemical vapor deposition has been used for the formation of solidfreeform objects.[1,2] The technology for Selective Area laser deposition or SALD and Selective area laser deposition and vapor infiltration, SALDVI can be modified for the purpose of joint formation and crack repair. In order to do this the crack or open space of the joint is defined as the inverse of a solid free form object and a computer scan path is generated for the laser. This technique has been used for the formation of silicon carbide joints of tubes in the SALD mode.[3,4,5] Joining attempts were made for flat plate specimens using the SALDVI method.[6] Silicon carbide and Silicon Nitride have been previously explored as the SALD joint fill material but oxides have not.

Joining of aluminum oxide using SALD or SALDVI has not previously been reported. Al2O3 deposited from the gas phase, is a material that is showing potential as a possible joint fill material, and precursors are being evaluated for this type of deposition.In order to deposit alumina onto alumina or onto another substrate with the SALD, orSALDVI process, a precursor must be selected. When performing a SiC deposition it is possible to utilize a precursor that, in one molecule, contains the Si and C needed to form the SiC. Tetramethylsilane (TMS) is such a precursor. Other ceramics will often require the use of multiple precursor chemicals. For instance, Si3N4 can be deposited with a combination of TMS and ammonia as a nitrogen source. In order to join alumina, a precursor must be selected that contains both oxygen and aluminum, or multiple precursors must be used.

In order to join alumina, a precursor must be selected that contains both oxygen and aluminum, or multiple precursors must be used. Besides this consideration, a viable precursor for the SALDVI process applied to joining should be one with a relatively high vapor pressure. This usually limits the selection candidates to a liquid or gas.For formation of alumina, several solid precursors exist and have been used in theformation of films but their use necessitates heating for vaporization. They are then primarily used in flowing system hot or cold walled CVD reactors.[7] For deposition using an unheated chamber with a fixed gas supply, this type of precursor is not suitable. Trimethylaluminum (TMA), (CH3)3Al was chosen as a source of aluminum because of its vapor pressure, 11 torr at room temperature; it is a liquid at room temperature.

Trimethylaluminum (TMA), (CH3)3Al was chosen as a source of aluminum because of its vapor pressure, 11 torr at room temperature; it is a liquid at room temperature. TMA has been used in several studies asthe Al source for alumina in certain types of chemical vapor deposition, particularly epitaxial monolayer growth. Although this type of growth is not suitable for the SALDVI process, the experiments indicate alumina is formed with this chemical precursor. The studies also indicate several oxidizer choices. Oxygen sources that have been used for Al2O3 growth include, N2O, NO2 H2O and H2O2.[8,9] Oxygen was not used in these studies because of the spontaneity of the reaction.TMA is pyrophoric alkyl and will react immediately with atmospheric oxygen if released. This necessitates special handling and experimental restrictions.

This necessitates special handling and experimental restrictions. In order to use this chemical for deposition studies, the vapor is extracted into a vacuum and the liquid is never exposed to atmosphere or directly to any process oxygen. The TMA is used only below its room temperature vapor pressure to prevent condensation on the chamber walls as a liquid. N2O or H2O were chosen as the first precursor possibilities. N2O as it is a non-toxic gas that can be easily introduced at any pressure. H2O was chosen because of its stoichiometric reaction potential: 3(H2O) + 2((CH3)3Al) ⇒Al2O3+ 6(CH4). Hydrogen is also added to this reaction though it is not stoichiometricly necessary. Hydrogen in excess, it is believed, assists in the formation of the desired CH4 product instead of creating hydrocarbon contamination.[10]Experimental Setup:The experimental setup for the formation of Al2O3 joint is a modified version of theSALD system.

A manifold of gasses, Argon, H2, N2O and O2 is attached to the system. The TMA and water are each uniquely valved off from the main system . The TMA feeds into a chamber that serves as a reservoir of vapor so that the TMA liquid does not contact the process gasses. There is a similar chamber for water vapor to prevent the TMA vapor from coming in contact with the liquid water. When introducing the precursors into the system it is desirable to introduce them in the vapor form. The sample work stage is within a vacuum chamber with a window for the introduction of the laser beam. The laser is moved externally by a computer controlled X-Y stage. The laser used in this study is a 1070nm wavelength fiber laser that is CW up to 100 watts. The laser is operated with a 70% beam planar beam splitter as there is instability below 10 watts. With the filter the laser is precisely controllable between 3 and 30 watts.

The laser is operated with a 70% beam planar beam splitter as there is instability below 10 watts. With the filter the laser is precisely controllable between 3 and 30 watts. The laser power and stage motion are controlled with a custom National Instruments Labview Code. This code is capable of interpolating a three dimensional machine path, either directly imputed or from a CAD file. The experimental setup is shown in figure 1.Figure 1: Schematic of Deposition and Joining Setup. A) Fiber laser 1070nm B)X-Y Stage with optical train C) Focusing Lens D) Reaction Chamber E) Gas Cylinders and Manifold F) Vacuum Pumps G) Liquid Precursor and Expansion Chamber H) Chemical Neutralizing BubblerThe experimental procedure for the determination of viable precursors is as follows. The reaction chamber and expansion chambers are flushed with argon and evacuated multiple times and a 50 mtorr or lower vacuum is established.

The reaction chamber and expansion chambers are flushed with argon and evacuated multiple times and a 50 mtorr or lower vacuum is established. The reaction chamber and one expansion chamber are filled with TMA vapor and then evacuated to one torr below vapor pressure, vapor pressure being approximately 11 Torr. If water is to be used as the oxidizer the second expansion chamber is filled with water vapor. The partial pressures within the expansion chambers are regulated to achieve the correct ratios upon mixing. Expansion chambers are used so that no liquid, H2O, H2O2, or TMA, can come in contact with a gas with which it might react. The liquids are thus mixed only in the gas phase. Once these precursors are mixed in the gas phase, additional process gasses are added. The gasses, which may be introduced into, the current set up include: hydrogen, argon, nitrous oxide, oxygen, and ammonia.

The gasses, which may be introduced into, the current set up include: hydrogen, argon, nitrous oxide, oxygen, and ammonia. Each gas is introduced through a manifold that has been evacuated so that each gas quantity is controlled.The purpose of these studies is to ascertain an oxidizer and aluminum source that can be used for pyrolytic laser CVD. A deposition is desired that will match the substrate composition, achieve a high degree of density, and show good adhesive properties to the substrate. Oxygen is not listed as a potential precursor to be used with TMA, since it reacts spontaneously. In order to test oxygen as a potential precursor, two torr O2 was introduced into a chamber with TMA at vapor pressure. Low oxygen partial pressure was tested for the purpose of finding a lower limit below which, a spontaneous reaction would be prevented or slowed to a negligible pace. Even at this low partial pressure the O2 rapidly reacted without any heat input from the laser.

Even at this low partial pressure the O2 rapidly reacted without any heat input from the laser. Below 2 torr the theoretical yield of Al2O3 would be to low. A fine alumina powder, which stayed suspended in the air for several minutes, was created by this reaction as well as potentially other byproducts. Studies indicate the reaction: [(CH3)3Al]2+O2⇒2(CH3)2AlOCH3 [11] may occur.Results:Water vapor and N2O have both been used as oxidizer sources with TMA. Water vaporand TMA will react to form amorphous Alumina, tested by EDS and X-ray diffraction, if the reaction occurs as a liquid based reaction. This was tested by using an over pressure of H2O and TMA, achieved by heating both liquids so that their vapor pressures would be doubled and allowing this super saturation to mix as a gas. The resulting reaction occurred both within the gas phase and on the surface of the reaction chamber, depositing Al2O3 on all cold surfaces and constriction points.

The resulting reaction occurred both within the gas phase and on the surface of the reaction chamber, depositing Al2O3 on all cold surfaces and constriction points. When the water and TMA were maintained 1 torr below their vapor pressures the spontaneous reaction either did not proceed or proceeded at a very slow rate.Nitrous oxide did not react spontaneously with the TMA at room temperature at any partial pressure of N2O. Gas combinations of TMA, hydrogen and a single oxidizer source, either N2O or H2O did not yield a controlled deposition. By combining N2O and H2O with hydrogen, argon, and TMA, both SALD and SALDVI deposits were then possible.Using TMA as the aluminum source, and the combined oxidizer phase of H2O and N2O,alumina deposits were formed, both as a deposition into a powder bed and a deposition onto a substrate both alumina. The SALDVI deposition occurred within an alumina powder bed.

The SALDVI deposition occurred within an alumina powder bed. The bed was created by repeated powder placement, agitation, and flattening by metal scraper. The powder surface was placed at the focal distance of the primary lens. A laser power of 15 watts, CW, was used to create a 3 mm square raster pattern with 50-micron spacing between raster lines. The focused spot diameter at the powder surface was 100 micron. The laser scan speed was 25 microns/s. The TMA and water vapor were used slightly below their vapor pressures. Argon, Hydrogen and N2O were all added at 150 torr. EDS performed on these deposits confirms aluminum and oxygen ratios of matrix and powder regions of deposit that closely approximate Al2O3. X-ray diffraction indicates alumina. The resulting deposition is seen below in the ESEM micrograph, Figure 2:Figure 2: 500x and 1000x ESEM micrographs of SALDVI deposition of alumina into an alumina powder bed.

The resulting deposition is seen below in the ESEM micrograph, Figure 2:Figure 2: 500x and 1000x ESEM micrographs of SALDVI deposition of alumina into an alumina powder bed. Note the particles imbedded within the deposited matrix.The SALDVI deposit above appeared dense and was approximately the expected deposition geometry. Metallographic examination is in progress.Deposition of the SALD type was also performed onto an alumina substrate. The conditions for deposition were the same as above except the raster pattern defined a 1cm x 2 mm scan area and the laser passed each spot 6 passes as apposed to four of powder bed deposition. A deposition was formed with a crisp interface as seen below in figure 3:Figure 3: SALD type deposition of alumina (right) on a sintered alumina substrate (left) at 470X magnificationThe deposition was a distinctly different from the substrate but failed to achieve the desired density. The deposition did show very good adhesion to the substrate.

The deposition did show very good adhesion to the substrate. X-ray diffraction patterns of the above substrate and deposition are shown below in figure 4 as well as an alumina pattern, which closely matches both materials.70006000500040003000Deposition Substrate al203 71-1684200010000010203040506070809010010002 theta (deg)Figure 4: X-ray diffraction patterns of substrate, deposition and corresponding pattern from card 71-1684 alumina.In an initial SALDVI joint attempt was performed that resulted in a deposition intopowder and adhesion to one work piece face. The deposition was composed of only two laser raster passes and consequently resulted in a non dense deposit. Seen below in figure 5 is a deposit on a joint wall.Figure 5: ESEM micrograph of SALDVI type deposit joint fill material. The light fluffy region comprises the deposited material and the larger crystalline shards are powder that has been infiltrated.

The light fluffy region comprises the deposited material and the larger crystalline shards are powder that has been infiltrated. The base substrate is not seen in this image.Summary:Alumina Deposits were successfully made by combining the precursors oftrimethylaluminum, nitrous oxide, hydrogen and water vapor. It was possible to use a computer defined laser path to make a deposit of the same geometry. The deposits were well adhered to the substrates. Shorter laser exposure time, resulting from fewer laser passes, resulted in deposits that were not dense. The density and adhesion increased with increasing laser dwell time.As a precursor for the aluminum component, TMA has potential. A combination of H2Oand N2O serves as an oxygen source, however, the exact nature of the reaction is not known. There are difficulties in densification and localization of the reaction to the substrate or powder surface; compositionally the alumina is as desired.

There are difficulties in densification and localization of the reaction to the substrate or powder surface; compositionally the alumina is as desired. As these conditions are improved, then joining will be possible with the SALDVI technique for alumina substrates. Work in the area of joining is ongoing which is taking advantage of this precursor selection evaluation.References:[1] P. Calvert, R. Crockett, “Chemical Solid Free-Form Fabrication: Making Shapes without Molds,” Chem. Mater. 1997, 9, 650-663[2] J. E. Crocker, H. Wei, L. L. Shaw and H. L. Marcus, “SALDVI of SiC into metal and ceramic powders,” Proceedings of the Solid Freeform Fabrication Symposium, The University of Texas at Austin (2001), pp. 163–169[3] S. L. Harrison, “Selective Area Laser Deposition (SALD) Joining of Silicon Carbide with Silicon Carbide Filler,” PhD Dissertation, University of Connecticut, 1999[4] S.L. Harrison, H. L.

GINATE-BASED RAPID PROTOTYPING SYSTEMP.J. BARTOLO, R. LAGOA, A. MENDESDepartment of Mechanical Engineering, School of Technology and Management Polytechnic Institute of Leiria Morro do Lena, Alto do Vieiro, Leiria, Portugal Reviewed, accepted August 13, 2003 AbstractAlginate hydrogels are an import class of ionic biopolymers for medical and biotechnological applications, produced by the combination of alginate with a proper cross-linking agent. During the gel formation, cross-links can be formed between the alginate chains and cationic species, changing the elastic behaviour of the material that controls the volume change phenomena of the gels. This paper proposes a new rapid prototyping system to produce three-dimensional alginate structures by extruding, layer-by-layer, a previously prepared solution of sodium alginate in water mixed with a solution of calcium chloride, both of known concentration.

The building process to obtain these gel structures is described, from a chemical point of view, and some obtained structures are shown. The alginate concentration over both the kinetics and accuracy effects of the process is investigated. The preliminary findings of this research work promise to open an exciting new area for medical applications.Keywords: Alginate, Hydrogels, Biomimetics, Medical Applications, Rapid prototyping.IntroductionThe engineering activity, a design by human process, is usually portrayed as a unique human process, relying on the intuition, experience and judgment of engineers and designers, to create a representation of objects meeting a set of requirements. However, there is also a design by nature (Bartolo and Bartolo, 2003) process, the process through which all the organisms dynamically adapt to its environment, performing its different functions, such as respiration, reproduction, sensing, movement, etc.

This design by nature is an evolutionary transformation process with a unique goal, the survival of each population of organisms through the correct adjustment to the environment. This kind of design process has been optimised throughout years and years of natural evolution.Just as an ecosystem incorporates change and adapts towards self-sustaining processes, lessons from nature can be used to support and optimise the design by human activity. The philosophy under this research project is the belief that engineers can learn with nature. This process of borrowing the best from nature is called biomimetics (Jeronimidis and Atkins, 1995) and this research study focuses on this new concept, investigating a new route to produce three-dimensional patterns in alginate hydrogels for medical applications, through a biomimetic free form manufacturing system.

This system intends to replicate natural procedures used by some marine brown algae, namely Laminaria Hyperborea, to produce alginate, a structural component of the algae.Hydrogels have been receiving much attention due to their potential use in a wide variety of biomedical applications, including tissue engineering scaffolds, drug delivery, contact lenses, corneal implants and wound dressing (Griffith and Naughton, 2002; Garcia and Ghaly, 1996;60Kim et al., 1998). Most hydrogels are usually synthesized using reactive monomers and cross- linkers. However, they do consume non-renewable resources, employ hazardous manufacturing steps and are not biodegradable, apart from not always being biocompatible so their use in medical applications is limited. Consequently, research attention has been centred on generating hydrogels from natural and biocompatible polymeric materials, such as collagen, gelatin, chitosan and alginate.

Consequently, research attention has been centred on generating hydrogels from natural and biocompatible polymeric materials, such as collagen, gelatin, chitosan and alginate. Biomaterials present properties, such as biocompatibility, biodegradability, sensitivity to pH, electrical or temperature stimuli, that enables its use in biotechnology and medical applications.Biologic manufacturing systemThe alginate-based rapid prototyping system, which is in a preliminary research stage, intends to produce alginate solid structures by extruding a previously prepared solution of sodium alginate in water mixed with a solution of calcium chloride, both of known concentration.

Understanding the reaction kinetics of the gelation process of alginate-based systems is the first step for the development of an intelligent system leading to the design, optimisation and control of an alginate-based rapid prototyping system.Alginate is a natural linear polysaccharide that contains 1,4-linked b-D-mannuronic (M) and a-L-guluronic (G) acid residues (Figure 1), arranged in a non-regular and block-wise fashion along the chain (Kang et al, 2002).

The biological properties of this biopolymer have been exploited in numerous medical and surgical applications, such as drug delivery, cell encapsulation, etc (Gombotz and Wee, 1998; Bucke, 1987).OHOHHOHOHHHCOO b D Mannuronic acida L Guluronic acidOHOHHOHOOFigure 1: Structure of an alginate showing a linkage between the mannuronic and guluronic acid.Gelling occurs when divalent ions (Ca2+, Ba2+, Fe2+, Sr2+, etc.) or trivalent ions (Al, etc.) take part in the interchain ionic binding between G-blocks in the polymer chain giving rise to a three dimensional network. Such binding zones between the G-blocks are often referred to as “egg boxes” (Figure 2). This ions act as cross-linkers that stabilise alginate chains forming a gel structure, which contains cross-linked chains interspersed with more freely movable chains that binds and entraps large quantities of water.

The gelling process is characterised by a re-organisation of the gel network accompanied by the expulsion of water (Serp et al., 2002).Its reactivity with calcium and the subsequent gel formation capacity is a direct function of the average chain length of the G blocks, showing a great affinity for Ca2+ ions (Gombotz and Wee, 1998). Gels made of M-rich alginate are softer and more fragile, and may also have lower porosity. This is due to the lower binding strength between the polymer chains and to61the higher flexibilities of the molecules. The gelling process is highly dependent upon diffusion of gelling ions into the polymer network.Ca2+Sol structureGel structureFigure 2: Alginate gelling results from reticulation of the chains by calcium ions.MaterialSodium alginate was purchased at Panreac (Barcelona, Spain). Calcium chloride was supplied by Carlo Erba (Milano, Italy). All solutions were prepared with pure water, with conductivity of 0.054 mS/cm.

Calcium chloride was supplied by Carlo Erba (Milano, Italy). All solutions were prepared with pure water, with conductivity of 0.054 mS/cm. Alginate solutions were prepared by addition of weighted portions of sodium alginate to measured volumes of water. Due to their high viscosity, these solutions were agitated by orbital shaking for three hours at 50 ºC to ensure good homogeneity. Calcium chloride solution 5% (w/v) was obtained dissolving the salt in water. This solution was diluted to obtain solutions containing different concentrations of calcium chloride.ResultsTo evaluate the kinetics of the gelling process solutions, containing different concentrations of alginate and different concentrations of calcium chloride (CaCl2), they were prepared and mixed at room temperature.

The effect of the alginate concentration is shown in Figure 3, which describes the weight loss as a function of time for solutions containing different amounts of alginate (1% and 2%) mixed with a solution of 5% CaCl2. Figure 4 illustrates the variation of weight loss as a function of time, for a solution containing 2% of alginate mixed with solutions containing different amounts of CaCl2. The experimental data (discrete values), was fitted using a sigmoidal equation. Fitting the experimental data enables us to have a continuous range of values to better evaluate the gelation process. To correlate the experimental data and the values obtained from the sigmoidal equation, a numerical routine using the Marquardt-Levenberg (Press et al., 1986) multivariable non-linear regression method was employed.

A good correlation was achieved by controlling the number of steps, the increment and a small tolerance parameter, corresponding to the difference of values at the step n+1 and values at the step n, which is used for convergence purposes.620,40,300,250,3Alg 2%Alg 1%0,200,20,150,100,10,050,00,00020406080100120140020406080100120Time (min)Time (min)Figure 3: Weight loss vs gelation time for two solutions containing different concentrations of alginate mixed with a solution of of 5% CaCl20,250,30CaCl2 1%0,200,25CaCl2 5%0,200,150,150,100,100,050,050,000,00020406080100120140020406080100120Time (min)Time (min)Figure 4: Weight loss vs gelation time for a solution containing 2% of alginate mixed with solution containing different concentrations CaCl2.The weight loss increases with time and is more significant for samples produced by solutions containing low contents of alginate and high contents of calcium chloride is shown in Figures 3 and 4.

This is due to two main reasons:ßthe amount of water present in the initial alginate solution, which is higher in more dilute solutionsßthe kinetics of the gelation process, which is higher whenever solutions containing higher concentrations of CaCl2 are used.The calcium divalent ions act as cross-linkers that stabilise alginate chains forming a gel structure. Therefore, increasing the concentration of calcium chloride present in the solution increases the cross-linking of polymeric chains and the expulsion of water. During the gel formation, calcium ions diffuse from the liquid solution to the forming gel and its complexation by the carboxylic groups of G blocks increases the gel stiffness. This phenomenon is accompanied by a significant loss in water at counter-current of the flux of calcium ions (Figure 5).

This phenomenon is accompanied by a significant loss in water at counter-current of the flux of calcium ions (Figure 5). Ahmad and Huglin (1994) and Hu and Lin (1994) claim that the water contained in alginate gels can been found in two states: free-water and non-free-water.63140140Therefore, it is expected that the water leaving the gel during the gelation process is the one which interacts less with the polymer (free-water).External environmentwaterAlginateCa2+Figure 5: The diffusion of chemical elements during the gelation process of alginate structures.Figure 6 compares the variation of the rate of weight loss with time as a function of both alginate and CaCl2 concentration. This Figure shows that maximum value of the rate of weight loss is almost independent of the concentration of alginate or calcium chloride and occurs very fast, in less than one minute.

However, the rate of weight loss is higher for samples produced with solutions containing higher concentrations of alginate or solutions containing higher concentrations of CaCl2.0,250,200,18)0,20a))0,16b)0,140,150,125% CaCl20,100,100,082% CaCl22% Alg0,060,053% Alg0,041% CaCl21% Alg0,02f0,000,0001234i0123Time (min.)Time (min.)Figure 6: Variation of the rate of weight loss vs time as a function of a) alginate concentration for samples produced through a solution containing 5% of CaCl2; and b) calcium chloride concentration for samples produced through a solution containing 2% of alginate.The variation of the maximum value of the rate of weight loss with the concentration of calcium chloride for samples produced using a solution of 2% of alginate was fitted using a sigmoidal equation (r2=1) as shown in Figure 7. This variation reveals two main regions.

This variation reveals two main regions. The first region corresponds to the gelation processes performed with solutions containing concentrations of CaCl2 lower than 3%. In this region, the concentration of CaCl2 has great impact on the rate of weight loss and consequently on the gelation kinetics. The second region appears at high values of CaCl2. This can be attributed to diffusion limitations of the divalent ions of Ca2+ as the gel structure is formed.

The second region appears at high values of CaCl2. This can be attributed to diffusion limitations of the divalent ions of Ca2+ as the gel structure is formed. Analogous behaviour is observed in Figure 8 for644650,420,440,400,380,400,360,360,340,320,320,300,280,280,260,24b)0,24a)0,220,200204060801001201400,100,150,200,25Time (min.)Weight loss (%)Figure 9: Variation of the shrinkage as a function of a) time and b) weight loss of a disc produced through a solution with 2% of alginate mixed with a solution containing 5% of CaCl2.The accuracy in the definition of the geometric shape of alginate structures is also strongly dependent on both the concentrations of alginate and calcium chloride. Figure 10 displays the shapes of different circular structures produced by mixing a solution with 2% of alginate with solutions containing different concentrations of CaCl2.

Figure 10 displays the shapes of different circular structures produced by mixing a solution with 2% of alginate with solutions containing different concentrations of CaCl2. A better definition is obtained using high contents of alginate and CaCl2. The gelation process of alginate structures will enable to retain the shape of structures during a large period of time (see Figure 11), owing to the fast process of he gelation rate during the first 5 minutes, that slows down afterwards slowing down the rate of weight loss.Figure 10: Circular structures produced by mixing solutions containing different concentrations of alginate with solutions containing different concentrations of CaCl2.66Figure 11: An alginate structure three weeks after its generation.A phenomenological model previously developed for stereo-thermo-lithographic processes (Bartolo and Mitchell, 2003; Bartolo and Mitchell, 2001) was adopted to evaluate the kinetics of the gelation process of alginate systems.

According to this model, the rate of weigh loss, dW/dt, is given by the following equation:dW dt⋅G=( m W1W)nwhere G is the gelation rate constant, W is the fractional value of weight loss, and m and n are the exponential coefficients of the kinetic model. All the kinetic parameters have been found to be dependent on both the alginate and calcium chloride concentration.The rate constant is given by the following equation:⋅=G xÊ ÁÁ exp Ëg [ CaCl2ˆ ⋅˜˜ ] ¯Ê ÁÁ exp Ëj [ lg Aˆ ]˜˜ ¯where x, g and j are constants, [CaCl2] is the concentration of calcium chloride and [Alg] is the concentration of alginate.

concentration of CaCl2 for samples produced using a solution containing 2% of alginate.The model intends to provide:ßa better understanding of the gelation process;ßa more accurate estimation of build times and weight loss profiles;ßa platform for testing new alginate structures or other bio-polymeric systems and process improvements;ßthe development of a standard set of parameters and gels for some pre-determined three-dimensional model characteristics.ConclusionsA new rapid prototyping system to produce three-dimensional alginate structures has been described from a chemical point of view. The gelation process is strongly dependent on both the alginate and calcium chloride concentrations. During the gelation process a significant water loss is observed influencing the shrinkage of the produced structures.

During the gelation process a significant water loss is observed influencing the shrinkage of the produced structures. Stable alginate structures are obtained after 20 minutes and the shape of these structures is retained for a long period of time.A kinetic model, describing the gelation process, is also proposed. This computational model will support the development and optimisation of this alginate-based rapid prototyping system, avoiding a great amount of experimental work.686Future work will focus on the relationship between morphology/mechanical properties and gelation kinetics in order to better understand the generation process of gel structures. This will help to produce appropriated porous structures for tissue engineering.ReferencesAhmad, M.B.

This will help to produce appropriated porous structures for tissue engineering.ReferencesAhmad, M.B. and Huglin, M.B., DSC studies on states of water in cross-linked poly(methyl methacrylate-co-N-vinyl-2-pyrrolidine) hydrogels, Polymer Int., 1994, 33, 273-277.Bartolo, H.M and Bartolo, P.J., Learning from biological systems, in Innovative Developments in Architecture, Engineering and Construction, Edited by C.J. Anumba, Millpress, Rotterdam, 2003.Bartolo, P.J. and Mitchel, G., Modelling the curing kinetics of thermosetting resins in stereolithographic applications, Proceedings of the 7th European Conference on Advanced Materials and Processes, Published by Associazione Italiana di Metallurgia, 2001.Bartolo, P.J. and Mitchell, G., Stereo-thermal-lithography: a new principle for rapid prototyping, Rapid Prototyping Journal, 2003, 9, 150-157.Bucke, C., Cell immobilization in Ca-alginate, Methods Enzymol., 1987, 135, 175-188. Garcia, A.M.

 ALGORITHM-BASED METHOD FOR PROCESS-SPECIFIC THREE- DIMENSIONAL NESTING FOR ADDITIVE MANUFACTURING PROCESSESAlexander Arndt*, Heike Hackbusch*, Reiner Anderl*Department of Computer Integrated Design (DiK), Technische Universität Darmstadt, 64287 Darmstadt, GermanyAbstractTo achieve optimal and high-quality results through additive manufacturing, the process- and technology-specific orientation and positioning of components within the virtual space, the so- called nesting, is essential. Primarily the nesting step is examined in this paper. From a scientific perspective it is a matter of examining this process and furthermore to analyze the optimal insertion of supporting structures, since the critical machine-specific parameters have been insufficiently studied. Within this paper a new multi-criteria optimization based on a conceptual algorithm is proposed. The most important point is the consideration of a technical and not only geometric nesting process.

Within this paper a new multi-criteria optimization based on a conceptual algorithm is proposed. The most important point is the consideration of a technical and not only geometric nesting process. The objective is the demonstration of restrictions and boundary conditions and a first developing for a new approach for the nesting process. As an example, the influence of the orientation of the spring rate is presented with a sample component here. Furthermore, there will be a prototype implementation and a short validation. Finally, a brief conclusion and an outlook is given.Introductionto conventional processes. The technological development enables manufacturing of simple prototypes, small batches, and individual products. Due to the renunciation of formative tools and material-intensive subtractive machining methods, costs can be lowered, while more particularly the additive manufacturing process chain can be shortened significantly, accelerating the process.

Real models and prototypes can be manufactured straight on the base of digital 3D CAD models. Thus, rapid manufacturing or the use of rapid prototyping, which takes advantage of the direct linkage to the CAD-interface, benefit the product development process.Additive manufacturing processesimply considerable advantages comparedThe additive manufacturing process chain characterizes the essential steps of the product development. The essential steps of the process chain are the preparation of the building process, the manufacturing of the physical model itself, followed by the post-processing and finally the use of the product. According to [Gib10], the process chain can be categorized into eight different steps, as can be seen in Figure 1. The first process step compromises the creation of the 3D CAD model, which contains all necessary geometric data regarding the constructing guidelines.

The first process step compromises the creation of the 3D CAD model, which contains all necessary geometric data regarding the constructing guidelines. The creation of the CAD model is followed by the conversion into a STL (Standard Tessellation Language) data format by triangulation. This step is a de facto industry standard. Given the STL data format, within the next step the parameters size, position and orientation of components are defined, the so-called nesting. Additionally, if needed, support structures are determined and generated, and a sliced model is passed on to the manufacturing machine. For the manufacturing, further steps as the choice of material are necessary. After the manufacturing, the last three stepsof the additive manufacturing process chain consist of the withdrawal of the product, the possible post-processing, and the utilization of the product.Figure 1: Process chain additive manufacturing.

Adapted from [Gib10]MotivationPrerequisite for an economically feasible utilization of any additive manufacturing process is a best possible preparation of the process. The process- and technology-specific orientation and positioning of components within the virtual space is essential in order to achieve optimal and high- quality results through additive manufacturing. From a scientific perspective this process still has potential to be analyzed and examined to achieve optimal solutions by using a conceptual algorithm. With an optimal nesting, which consists of the positioning and orientation of components, negative effects as the staircase effect, curling, surface roughness or component accuracy can be enhanced significantly.State of the ArtAs current research shows, the nesting process and the filling of the manufacturing work space can be categorized into different bin packing problems.

The classification by Dyckhoff [Dyc90] is used for the identification of the arrangement and categorizes the dimension, the distribution, the room assortment, and the range of goods. All problems are NP-complete (non- deterministic polynomial-time-complete) and can be approximated with genetic algorithms. As Table 1 gives an overview, the main bin packing problems are the knapsack problem, the container loading problem, the pallet loading problem, and the general cutting stock problem. All algorithmsattempt an optimal filling of the virtual space based on the geometric data, packing density and the orientation of the components.Table 1: Overview of the main bin packing problemsPacking problemDescriptionFigurem e l b o r pk c a s p a n KThe knapsack problem can be illustrated visually with the optimum filling of a knapsack. The knapsack has a limited volume and every object which is to be packed has a specific volume and value, e.g. the weight.

The knapsack has a limited volume and every object which is to be packed has a specific volume and value, e.g. the weight. The aim of optimization should be to load the knapsack with the maximum weight but without exceeding the volume of the knapsack with the sum of the object volumes. [Wäs07]g n i d a o lt e l l a pm e l b o r pIf rectangular block objects of the same size are given and to be placed on a rectangular base while achieving the maximum number of objects accommodated without exceeding the maximum, one speaks of pallet loading problem. [Sch08b]g n i d a o lr e n i a t n o cm e l b o r pThe container loading problem can be illustrated visually by packing a container with rectangular block objects. Compared to the pallet loading problem, the objects have different dimensions. [Sch08b]k c o t sg n i t t u cm e l b o r pThe cutting stock problem can be presented pictorially.

Compared to the pallet loading problem, the objects have different dimensions. [Sch08b]k c o t sg n i t t u cm e l b o r pThe cutting stock problem can be presented pictorially. There is a defined number of objects with defined volumes available to be distributed to as few knapsacks as possible. The knapsacks have a maximum capacity. [Sch08b]First research specifically for additive manufacturing was done by Corcoran [Cor92]. With a 3D Next Fit Algorithm and a 3D First Fit Algorithm cuboids are placed into a virtual space. An advanced algorithm was applied by Ikonen [Iko97] to place nonconvex, three-dimensional figures with cavities in a virtual space. It was optimized for selective laser sintering. The component cavities are used for positioning but component penetration has to be eliminated manually and theinfluence of orientation on the component quality is disregarded.

The component cavities are used for positioning but component penetration has to be eliminated manually and theinfluence of orientation on the component quality is disregarded. Hence, Lutters, Dam and Faneker used the “brazil nut”-effect to simulate three-dimensional nesting of complex figures [Lut12]. Based on the criteria “Preferred Nesting Orientation” and “Preferred Quality Orientation” it is possible to exclude component penetration, but optimal utilization of the virtual space cannot be reached and cavities are not used. For stereolithography an efficient nesting was developed by Canellidis [Can13]. The component positioning is done two-dimensionally due to the negative effect of stacking in the stereolithography process. The component orientation is assumed to be ideal and therefore, not changed during the packing. To determine the packing, different deterministic approaches as the Left-Bottom/Down-Bottom and the No-Fit Polygon are considered.

To determine the packing, different deterministic approaches as the Left-Bottom/Down-Bottom and the No-Fit Polygon are considered. However, in all algorithms solely geometric parameters are considered which are not machinery-specific yet utilization-oriented. Hence, an analysis in its entirety with all manufacturing and component restrictions is necessary, see also Figure 2. The current geometric STL data format is irreversibly faceted as a result of the staircase effect, surface roughness and other effects when transforming and approximating the CAD data. To achieve optimal nesting which is not only based on geometric data, but also considers further process parameters, a new data format and nesting process has to be developed.

To achieve optimal nesting which is not only based on geometric data, but also considers further process parameters, a new data format and nesting process has to be developed. Ideally, a consistent, general data format which processes geometric and orientation data as the STL format [Arn14], additionally considers and gathers information concerning the material properties, endurance strength of the material based on the chosen orientation in the nesting process and post-processing. A consistent information and data management enables an optimal nesting process, but also optimal component quality, support and economic viability. In the first step, the STL format will be further used and build on the nesting investigated.

In the first step, the STL format will be further used and build on the nesting investigated. With a multi-criterial optimization of the process, an algorithm processing all the different information and boundary conditions has to be developed.Boundary Conditions and RestrictionsThe nesting process has a significant influence on the additive manufacturing process as well as on the final product. Therefore, it has to be well considered when establishing a consistent information and data management with information additionally concerning the material and the manufacturing process. Boundary conditions and restrictions concerning the geometry are explained in the following, but in order to accumulate restrictions in its entirety, boundary conditions concerning the manufacturing process have to be considered furthermore. The aim is to achieve a technological- and process-specific nesting process rather than just the common geometrical nesting.Figure 2: geometrical nesting vs.

The aim is to achieve a technological- and process-specific nesting process rather than just the common geometrical nesting.Figure 2: geometrical nesting vs. process-specific nestingAll additively manufactured components use a construction based on successive layers of material. After the slicing process of the STL data, the component is manufactured. The resulting difference in the contour of the single layers is the so-called staircase effect. The effect depends on the density, the surface angle of the mounting direction and influences the surface finish as well as the fidelity of transformation. The thinner the layers and the higher the number of layers, the higher the quality of the surface and the smaller the staircase effect [Bre13]. These positive factors are in contrast to the process time and therefore, the cost of the process. Another factor influencing the cost is the use of support structures.

These positive factors are in contrast to the process time and therefore, the cost of the process. Another factor influencing the cost is the use of support structures. Some processes need additional support structures to fix the manufactured component in place. Support structures require elaborate post-processing as the depletion of the additional support material and surface finishing [Gur14]. According to [Mar11], it is additionally important to control the temperature gradient during the manufacturing process in order to avoid curling and negative effects on the mechanical properties. Temperature gradients may lead to a reduction of the surface quality and therefore, are to be avoided. Nevertheless, an advantage of the additive manufacturing process are the scarcely restrictions concerning the manufacturable geometries of the components [Ver14].

Nevertheless, an advantage of the additive manufacturing process are the scarcely restrictions concerning the manufacturable geometries of the components [Ver14]. Indentation and cavities are producible, which enhance the ability of an optimal packing of the virtual space.The influence of the nesting on the final product can be categorized into the resulting strength of the material, the surface quality, the curling and the wall thickness [Weg12]. An anisotropy in the mechanical behavior and a lowered ultimate tensile strength of the component always result depending on the orientation and positioning of the component. Kirchner [Kir10]tested different tensile samples in terms of ultimate tensile strength and ultimate strain with results pending ±25% around the average.

Kirchner [Kir10]tested different tensile samples in terms of ultimate tensile strength and ultimate strain with results pending ±25% around the average. Best results in terms of tensile strength could be accomplished with a horizontal manufactured component, which achieves twice the ultimate tensile strength compared to vertical manufactured samples according to Schäfer [Sch08a]. Research still needs to be done to integrate the strength into a genetic algorithm in order to regard the mechanical behavior already during the nesting process. As mentioned earlier, the surface finish also depends on the positioning and orientation of the manufactured component. Depending on the depletion of support material and the staircase effect, the roughness average and roughness depth differs significantly with 200µm according to [Sch08a].

Depending on the depletion of support material and the staircase effect, the roughness average and roughness depth differs significantly with 200µm according to [Sch08a]. Aftertreatment can improve the surface finish, but is always linked to a deviation of the geometry.Figure 3 gives an overview of the interactions between nesting-dependent component and process characteristics. For this paper, these rudimentary descriptions should suffice. Since the physics of each process are so complex and different only worked with assumptions is below. For a detailed description of the physical quantities and effects, reference is made to the relevant literature.Figure 3: Interactions between a nesting-dependent component and process characteristics.

Adapted from [Dan10]Based on the shown packing problems and the boundary conditions and restrictions, a process-specific three-dimensional nesting for additive manufacturing processes is derived in the following conceptual design.ConceptAs it can be seen in Figure 2, the nesting is to be conducted with regard to technological- and process-specific characteristics. With respect to the preceded research, the state of the art, and the requirements for the manufacturing, the following requirements for the concept of the process- specific three-dimensional nesting have to be fulfilled. Simultaneously, an objective function has to be defined.The optimization is divided into two individual optimizations with its own genetic algorithm, as can be seen in Figure 4. In the orientation optimization a limited number of optimalx-y-orientations is determined for the components (STL files). Each of these x-y-orientations is then rotated in x-steps around the z-axis.

Each of these x-y-orientations is then rotated in x-steps around the z-axis. In every step, a voxel is generated. In this voxel, the part geometry, the orientation and the distance between components to be observed is displayed. In a second optimization step, the optimization of the component positioning, the voxel is to be placed inside the virtual space then tried by each component.Figure 4: algorithm of optimizationOptimization of the component orientationThe orientation of the component in the virtual space in relation to the orientation of the manufacturing is essential for the processing of the component and has a significant influence on the achievable manufacturing quality and the characteristics of the manufactured component. In general, it is necessary to consider the geometric target variables as the optimal packing height, the volume utilization, as well as the capacity utilization.

In general, it is necessary to consider the geometric target variables as the optimal packing height, the volume utilization, as well as the capacity utilization. Additionally, for the process-specific nesting, the curling effect should not be neglected and has to be implemented into the packing-algorithm. The effect is solely referring to shape deviation and can be categorized in four ways in terms of the additive manufacturing process, which implements a machinery-specific consideration: thermal curling, mechanical curling, the staircase effect, and component and endurance strength.Based on this knowledge, the curling effects are used as optimization criterions for the objective function (cid:1858)(cid:3042)(cid:3045)(cid:3036)(cid:3032)(cid:3041)(cid:3047) of the orientation, see equation (1).

Even though the computed area of the error triangles can differ slightly from the actual error, this computation seemed to achieve reasonable estimations in first evaluations. The smaller the result of criteria Kritopt,sce more better optimization exists. This also applies to the following criteria.In processes that produce material bond thermally or by means of photopolymerization, may lead to thermal curling due to the heat input. The optimization criteria (cid:1837)(cid:1870)(cid:1861)(cid:1872)(cid:3042)(cid:3043)(cid:3047),(cid:3047)(cid:3030) for the thermal curling resulting from temperature gradients is not fully understood yet and therefore, is based on empirical values according to [Dan10].

Since the heat input with the exposed surface contour of a layer is related to the area can use to detect the effect, see [Dan10].The curling effect, called mechanical curling, as a result of additive manufacturing processes using a bed of powder, is represented by the factor (cid:1837)(cid:1870)(cid:1861)(cid:1872)(cid:3042)(cid:3043)(cid:3047),(cid:3040)(cid:3030). The bigger the cross section of the first manufactured layer and the more weight is put on it, the bigger the effect will be. Since all layers are based on the bottom layer, the mechanical curling increases over the height of the structure, influencing the shape deviation. Hence, the criteria (cid:1837)(cid:1870)(cid:1861)(cid:1872)(cid:3042)(cid:3043)(cid:3047),(cid:3040)(cid:3030) is the area of the cross section in the orientation relative to the volume of the component.Furthermore the component and endurance strength is contained with acceptance of conditions in the objective function.

Here is an example to explain the spring rate k. Tests have shown that different spring rates in different orientations were present, see Figure 5. For this, samples were manufactured by Fused Deposition Modeling and subjected tension-pressure- experiments. This experiments have shown, if the component is manufactured lying, the spring rate higher (k=2,18N/mm) than the standing component (k=1,23N/mm). This physical effect is due to the anisotropy of the material or component. These scientific findings are integrated as constant values in the objective function (cid:1858)(cid:3042)(cid:3045)(cid:3036)(cid:3032)(cid:3041)(cid:3047). The minimization of the objective function is carried out to the optimum of the highest possible spring rates.

The minimization of the objective function is carried out to the optimum of the highest possible spring rates. The higher the spring rate, the lower the optimization criteria Kritopt,cs.Figure 5: Influence of the orientation of the spring rate (left: example components, middle: test preparation, right: diagram with results for spring rates)Optimization of the component positioningAs described previously, not only the orientation but also the positioning of the components in the virtual space can have a significant effect on the manufacturing process and the component’s quality. The better placed, the less time will be needed to manufacture it and therefore, the more economically feasible it will be.

The number of components solely indicates how many components can be manufactured in one process and therefore, are a measure for the number of client’s order can be accomplished ((cid:1837)(cid:1870)(cid:1861)(cid:1872)(cid:3042)(cid:3043)(cid:3047),(cid:3041)(cid:3030)(cid:4667). This criteria is calculated by dividing the components housed in the pack and all components to be manufactured. This result is always less than or equal to 1, therefore, it is withdrawn in the objective function of 1 (1 - Kritopt,nc). The greater is the result of this division, the better is the optimization result.Representing the packing height, the factor (cid:1837)(cid:1870)(cid:1861)(cid:1872)(cid:3042)(cid:3043)(cid:3047),(cid:3043)(cid:3035) gives an approximation of the needed process time, which depends significantly on the number of layers and is proportional to the packing height.

These are the subsystems implemented optimizations of the component orientation, optimization of the component positioning, and the user interface. The visualization of the program flow is shown in Figure 6. First, the STL files are loaded into the implementation (step 1). In the next step, the optimal x- and y-orientations are calculated by means of the genetic algorithm (step 2). Thereafter, the components about the z-axis are rotated and the optimum distances are calculated (step 3). This is done for all loaded STL files in the implementation. The calculation are then followed by the start of the optimization of the positioning of the components (step 4). When the optimization is completed, the best package is visualized. The result is the nesting as an STL file. Then support generating and the slicing can occur.In the steps of the optimization of the orientation and the positioning come genetic algorithms used.

The result is the nesting as an STL file. Then support generating and the slicing can occur.In the steps of the optimization of the orientation and the positioning come genetic algorithms used. The sequence of these algorithms is to look in the Figure 6 above right. For a better understanding now a digression on the subject genetic algorithm. In a genetic optimization a solution candidate is called individual. In analogy to biology is called an amount of solution candidates population. The genetic algorithm developed a population using evolutionary operators (recombination and mutation). This period is so long until a termination condition is met. The fittest individuals survive or form the basis for new individuals (selection). The fitness is determined by an objective function. For detailed information about genetic algorithms see specialist literature.The optimization of the component orientation is divided here into two parts.

For detailed information about genetic algorithms see specialist literature.The optimization of the component orientation is divided here into two parts. Both parts, the x-y-orientation and the z-orientation work with the genetic algorithm. After loading the STL Files and the input of the parameters starts the optimization. Here the objective function on the genetic algorithm is minimized. By means of selection, recombination and mutation are found the fittest orientations for the respective components. If a predetermined termination condition is satisfied, the algorithm is terminated, this is not met further optimized, see also Figure 6 the upper right corner. If all components are optimally orientated start the optimizing of the positioning. Analogously to the first optimization a genetic algorithm is used here also. Here are necessary as input the file paths of the components as well as the results from the first optimization.

Analogously to the first optimization a genetic algorithm is used here also. Here are necessary as input the file paths of the components as well as the results from the first optimization. Result is an array that contains all analyzed in the course of evolution component packages with associated indicative criteria, objective function values and matrices. About a function, the positioncoordinates and the angle of orientation of the components are issued. Based on this data an STL file is created, which visualizes the optimized pack.Figure 6: prototype implementation - programm flow. Adapted from [Fri14]To validate serves the example of optimizing of the component positioning .To check the package optimization for plausibility the results of isolated optimization processes are considered and compared to individual target criteria. For example: in a cylindrical space twelve parts should be packed. However, this can not be all housed together due to their geometry.

For example: in a cylindrical space twelve parts should be packed. However, this can not be all housed together due to their geometry. The algorithm should now solve succession with four different objectives, the packing problem.1. As many parts to accommodate (Kritopt,nc) 2. As compact pack (Kritopt,cmpct) 3. As much of the total volume of the building space used (Kritopt,vol) 4. The packing height minimize (Kritopt,ph) The weighting factors of the remaining optimization criteria have been set in this validationto zero. Figure 7 shows the visualization of the four different objectives.Figure 7: Component packages with different objectives. Left: side view in the y-direction; right: top view. Adapted from [Fri14]The analysis of the algorithm shows that most of the parts (10) could be accommodated with alignment on the optimization criteria Kritopt,nc in the installation space. Similarly many components adapted to optimize the use of space Kritopt,vol.

Similarly many components adapted to optimize the use of space Kritopt,vol. This can be explained by the compactness of the components. These have only barely cavities. It should be noted, that was terminated at this plausibility checks the genetic algorithm after a relatively few generations. The results are thus to be seen only as a trend. It is to complete a detailed validation.Conclusion and OutlookThe prototypical implementation confirms the function of the developed concept. The concept considers the optimization of the packing height, the volume utilization and the post- processes. Further, the consideration of the process-specific values is implemented.This paper presents a rudimentary concept for optimizing the part orientation and arrangement of several components for additive manufacturing processes.

At this point, pending steps and some development potential of this concept should be demonstrated.Within a next step, the developed mechanisms and introduced characteristic values should be thorough validated.Only certain effects which occur in additive manufacturing processes, e.g. the staircase effect, are considered in the here presented basic concept. Hence, in subsequent research a focus has to be put on extending the objective function of the genetic algorithm by the missing effects.For a user friendly field introduction of the algorithm, the profiles of the parameters are to be constructed on the base of the necessary machinery specifications.To improve the performance of the algorithm, it might be useful to implement it in a hardware-compliant programming language. Within the first parametrization of the two algorithms, standard values found in the literature are considered.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAligning Material Extrusion Direction with Mechanical Stress via 5-Axis Tool PathsJ. A. Gardner*, T. Nethercott-Garabet*, N. Kaill*, R. I. Campbell, G. A. Bingham*, D. S. Engstrøm^, and N.O. Balc+Loughborough Design School, Loughborough University, United Kingdom ^Wolfson School of Engineering, Loughborough University, United Kingdom +Department of Manufacturing Engineering, Technical University of Cluj-Napoca, RomaniaAbstractMechanical properties of parts fabricated via the Material Extrusion (ME) process can be improved by optimising process settings, however, their properties are strongly influenced by build orientation due to the stair-stepping effect initiating cracks whilst under load.

5-axis ME enables the fabrication of parts without the layer-by-layer restrictions that conventional 3-axis strategies impose. By aligning extrusion direction with high stress tensors, 5-axis tool paths can be used to reduce the effects of weak inter-layer bonds.To establish performance differences between parts manufactured by either strategy, wave spring-inspired geometry was selected for production, due to the multi-directional tensile loads acting throughout the material. 5-axis and 3-axis tool paths were generated via the Grasshopper 3D virtual environment within Rhinoceros 3D and MakerBot Desktop, and manufactured using a 5AXISMAKER and a MakerBot Replicator 2, respectively.

To evaluate performance differences between the two strategies, compression tests were conducted on the parts.Keywords: Additive Manufacturing, 5-axis, Tool path, Material Extrusion, G-codeIntroductionMaterial Extrusion (ME) is an Additive Manufacturing (AM) process category in which material (usually a thermoplastic polymer) is heated and extruded onto a substrate. The semi- molten filament forms a partial bond with previously deposited material, cools, and solidifies to form a layer of material [1]. The unique capability of AM typically relies on the translation of Computer-Aided Design (CAD) geometry into Two-Dimensional (2D) cross-sectional areas which are then successively fabricated, layer-by-layer to produce near net shape parts. By adopting a layer-by-layer strategy, complex, multi-functional parts can be fabricated, minimising the waste, tooling, and fixtures associated with conventional manufacturing methods [2].

By adopting a layer-by-layer strategy, complex, multi-functional parts can be fabricated, minimising the waste, tooling, and fixtures associated with conventional manufacturing methods [2]. Despite these advantages, there are fundamental AM process limitations preventing its wide-scale adoption across manufacturing industry. Many of these limitations occur from the build-up of layers, namely: loss of dimensional accuracy, stair-stepping effect (leading to poor surface finish), and anisotropic mechanical properties. These limitations are especially problematic within ME for the manufacture of thin-walled, or curved surfaces [3]. Mechanical instability within ME parts is strongly influenced by reduced bond strength between successive layers. Also, because the polymer is only partially melted, full part density often cannot be achieved [4].

Also, because the polymer is only partially melted, full part density often cannot be achieved [4]. Additionally, the rapid heating of the polymer through the nozzle followed by cooling causes a2005phase change in the material resulting in volumetric shrinkage of the filament bead. This can cause warpage, internal stress, and porosity between filaments, thus reducing the strength of a part [5, 6]. ME parts can be described as an anisotropic composite of discontinuous polymer fibres and voids [7]. Within the context of producing complex functional end-use parts, ME is caught in a paradoxical scenario in which layer-by-layer fabrication is both its greatest enabler and limiting factor.The objective of this research was to investigate a novel AM process based on ME which uses 5-axis Computer Numerical Control (CNC) tool paths to deposit material in a non-planar strategy.

It was envisioned that altering the orthotropic accumulation of layers to a dynamic 3D accumulation, parts would exhibit: more consistent material properties, greater surface contact area between each ‘layer’, and improved surface finish, thus improving mechanical performance and predictability when compared with 3-axis built equivalents. This process could enable the fabrication of low cost, customisable, parts with properties suitable for end use, load-bearing applications.Several 5-axis controlled AM processes are commercially available, for example, they are often used within direct energy deposition; a form of laser welding. Whereas conventional 3- axis AM strategies slice geometry into simplified 2D tool paths, 5-axis tool path generation becomes dynamic, requiring specialist software not usually associated with AM. Curved layer Material Extrusion (CLME) is an AM process which has been developed over the past decade and presented by a number of authors [3, 4, 8].

Curved layer Material Extrusion (CLME) is an AM process which has been developed over the past decade and presented by a number of authors [3, 4, 8]. A ME machine is typically modified with an additional axis of rotational movement in the substrate allowing the tool movement required to deposit curved layers. Singameni et al. [7] compared the mechanical properties of ABS parts manufactured using regular 3-axis material extrusion and 4-axis CLME. It was found that curved-layer parts had a superior mechanical performance. The average maximum load until failure increased by 40% during a non-standard bending test. Surface finish of the curved layer parts was also improved since no ‘stair-stepping’ effect was observed, reducing the chance of cracks initiating at surface irregularities.

Surface finish of the curved layer parts was also improved since no ‘stair-stepping’ effect was observed, reducing the chance of cracks initiating at surface irregularities. Additionally, it was observed that the curved layers prevented crack propagation, delaying catastrophic failure of the samples.MethodTo compare the mechanical properties between parts built with 5-axis and 3-axis machine configurations, wave spring-inspired geometries, hereon referred to as ‘wave springs’ for simplicity, were fabricated in PLA. Non-standard mechanical compression tests were then conducted on each sample.For a wave spring in axial compression, an estimation of deflection (Eq.1.) and operating stress (Eq.2.) can be calculated using wave spring formula, where: P = load, E = tensile modulus of material, K = multiple wave factor, Dm = mean diameter of spring, Z = number of layers, b = Radial width, t = thickness, and, N = number of waves p/layer.

See Figure 1.𝐷𝑒𝑓𝑙𝑒𝑐𝑡𝑖𝑜𝑛 = 𝑓 =3𝑍 𝑃𝐾𝐷𝑚 𝐸𝑏𝑡3𝑁4𝐼.𝐷𝑂.𝐷, 𝐸 =3𝑍 𝑃𝐾𝐷𝑚 𝑓𝑏𝑡3𝑁4𝐼.𝐷𝑂.𝐷(1)𝑂𝑝𝑒𝑟𝑎𝑡𝑖𝑛𝑔 𝑆𝑡𝑟𝑒𝑠𝑠 = 𝜎𝑜𝑝 =3𝜋𝑃𝐷𝑚 4𝑏𝑡2𝑁2(2)2006Figure 1 - Wave spring geometryIt is widely reported that build orientation influences elastic modulus and ultimate tensile strength of ME materials [9, 10]. In a material specification sheet for PLA, Stratasys report a 20% reduction in the value of the elastic modulus and a 50% reduction in ultimate tensile stress depending on build orientation [11].The wave spring formula assumes the spring has homogeneous, isotropic material properties and therefore cannot be accurately applied to 3-axis ME parts since the part could fail at a considerably lower load, due to inter-layer shear exceeding the ultimate shear strength of the material.

It was envisioned that building up 3D shells, rather than ‘2.5’ layers, to achieve a final geometry would increase the maximum load of the spring in compression when compared to samples fabricated via conventional 3-axis AM. Additionally, as no inter-layer shear would occur the wave spring formula could be applied.5-Axis G-code Considerations and GenerationME of a Computer-Aided Design (CAD) model requires software-based processes to generate commands, or G-code, which are translated by a CNC machine controller into machine movement. Conventional ME machines operate in X, Y, and Z axes and typically build parts via series of one or two axes movements. An interpretation of 3D volumes as multiple, stacked 2D areas allows for the discrete, layer-by-layer generation of G-Code.The planar nature of 3-axis ME allows for many of the associated G-code generation processes to be automated, therefore requiring minimal user input.

Volumetric meshes are imported into virtual environments, process settings are selected, and 3-axis G-code is generated without further user involvement. The authors are unaware of any currently available software able to generate G-code that describes a 5-axis tool path immediately following a mesh import2007process. A new process order was therefore required. It was proposed that the initial creation of a tool path could precede the generation of G-code for 5-axis ME.In all forms of ME, parallel tracks are essential in limiting the variation of surface contact areas between adjacent tracks, as deviations from uniformity may cause bonding issues. Adjacent tracks that are too distant share reduced surface contact areas, whereas tracks too close to one another may cause nozzle blockages due to increased pressures within the extruder assembly. To avoid these problems distances between adjacent tracks were made constant; therefore track sections were designed to be parallel.

To avoid these problems distances between adjacent tracks were made constant; therefore track sections were designed to be parallel. 2D and 3D parallel tracks are shown in Figure 2a and Figure 2b, respectively.ParallelTranslatedabFigure 2 - A comparison of (a) a parallel curve and a translated curve, with equal-length distance indicators, and (b) an example 3D curve created by the developed component clusterFigure 2a shows that larger angles between the curve’s translation direction and its normal vectors increases the distance between the translated and parallel curves.

This is a geometry-based consideration rather than a direct G-code-based consideration; however, it was necessary to account for this whilst creating the 5-axis tool paths to reduce the chance of a build failure.5-axis ME requires new G-code generation methods to: (i) orientate a nozzle to be perpendicular to its travel direction, and; (ii) offset the nozzle tip from obstacles such as jigs and previously extruded material. Additional challenges are also present, such as extruding low- viscosity material through a nozzle not aligned with gravitational forces, and the addition of cooling fans increasing the potential for machine-part collision. These challenges can be addressed by the optimisation of process parameters and by improving system design.The reference position of a 3-axis ME system can be located at a nozzle tip centre (NTC) as there are no rotating axes to calculate for.

Less logic is required to define a reference position that is not affected by rotation, in comparison to defining a position that is affected by rotation. Therefore, 5-axis G-code generation includes constant offsets in the B and C axes to calculate coordinate values for X, Y, and Z.The 5-axis hybrid manufacturing machine, 5AXISMAKER, was controlled by Mach3 software [ 12 ], a G-code reader and executer. To generate the commands an open-source2008algorithm [13] was used in Grasshopper 3D [14] within Rhinoceros 3D software [15]. To create the points that the NTC of the 5-axis machine moved between, the open-source algorithm required a single reference curve to describe all its positions.The creation of the curve, or tool path, was completed via the following steps.

A series of points that described a circle were offset from a plane, whereby each point was independently translated by referencing: its index position in the series, a sinusoidal function, the selected wave amplitude, and the selected number of peaks in a layer, all within a component cluster developed by the authors (shown in Figure 3). These translated points were interpolated to create a closed sinusoidal waveform. This curve was subsequently recreated to form a set of parallel curves, which defined a complete wave spring layer. The distance between adjacent curves was calculated by dividing the layer thickness by the number of tracks in the layer. To form a complete shell, copies of the tracks were rotated and translated, according to the selected: number of layers, number of peaks, wave amplitude, offset between layers, and layer thickness (also shown in Figure 3).

The shell was then non-uniformly scaled with increasingly larger diameters according to the selected number of shells and shell gap.First shell Last shellTrack connectorsFigure 3 - The developed component cluster with a corresponding tool path showing the relationships between parameter input and output valuesAt this stage of the G-code generation process all curves, or tracks, were disconnected and required connecting to form a single tool path. To determine how each curve was connected to its neighbour, the 5-axis machine setup was considered. The cables that connected the electronics box to the extruder assembly were located externally from the central column; therefore it was important to consider potential cable entanglement issues.

The cables that connected the electronics box to the extruder assembly were located externally from the central column; therefore it was important to consider potential cable entanglement issues. To minimise the probability of entanglement issues and to reduce build time, the completion of each track was2009followed by the subsequent track being followed in the reversed direction, as shown schematically in Figure 4.Figure 4 - A top view schematic of the tool path directionAs multiple bimodal parameters were used to determine the direction of the central column, it was important to understand the relationships between the parameters and the directions of each track. The parameters were: the number of tracks in a layer, the number of layers in a shell, and the number of shells in a part (as shown in Figure 3).

A flowchart (Figure 5) also displays the top-level logic that was followed to determine track direction.YesNoStart each layer in alternating directions.Start each shell in the same direction.End (at bottom)End (at top)Even number of tracks per layer?YesEven number of shells?YesNoEven number of tracks per layer?Start each layer in the same direction.Start each shell in the same direction.YesNoStart each layer in the same direction.Start each shell in the same direction.Start(at bottom)Even number of layers per shell?NoStart each layer in alternating directions.Start each shell in alternating directions.Figure 5 - A flowchart to determine tool path direction to avoid cable entanglement of the 5-axis machine2010From the initial movement-based command of 5-axis G-code a sinusoidal wave was built onto a cylindrical jig.

Once the first wave was built the nozzle translated in the Z direction only, and then completed the subsequent track in the reversed direction. The number of tracks per layer and the number of layers per shell were calculated by the values of the input variables of the developed component cluster. The values of these bimodal parameters therefore determined the starting direction of each successive layer and shell. The directions of alternating tracks were then reversed, according to logic described in Figure 5, and straight line connectors (shown in Figure 3) were added to connect the end of one track to the beginning of the next. The purpose of this process was to combine a series of tracks into a continious track, allowing the open-source algorithm to calculate all lines of G-code.

The purpose of this process was to combine a series of tracks into a continious track, allowing the open-source algorithm to calculate all lines of G-code. An example of a complete tool path track can be seen in Figure 6.abcFigure 6 - (a) Perspective, (b) top, and (c) front views of a single curve that represents a wave springThe tool path curve was divided into equal length segments whose end points represented the positions of the NTC. Nozzle orientation was determined using the normal vectors of an open cylinder as a reference object, to calculate values for the B and C axes. The closest point on the reference object from each NTC position was calculated to select the normal vector that most closely aligned with the intended nozzle orientation. The outcome of this process therefore defined X, Y, Z, B, and C axis values and complete extruder assembly pose, whereby pose encompasses position and orientation.

The outcome of this process therefore defined X, Y, Z, B, and C axis values and complete extruder assembly pose, whereby pose encompasses position and orientation. Additional parameters were also incorporated into the G- code generation process, such as: extruder length, B and C offsets from the central column, and the number of equidistant points on the tool path.2011The G-code was exported to text file format. As the same logic can be used to generate many wave spring geometries via 5-axis AM, only the parameter values of the developed component cluster (shown in Figure 3) would need to be modified. Each set of values can create wave springs with significantly different characteristics, which would be suitable for a range of loading conditions.5-Axis ME Process DescriptionThe 5-axis ME process involved a system with process settings that incorporate angles of rotation in addition to 3-axis control, it can also be referred to as 5-axis curved layer deposition.

This technique allows for further improvements to the mechanical properties of AM parts by curving the layers to maintain a continuous track aligned to stress tensors, which should reduce shearing effects between adjacent track sections. It is important to note that this technique is still in its infancy with few machines capable of manufacturing in this manner (7,16).Prior to 5-axis part manufacture a substrate was prepared in the form of a machined aluminium jig. Material was then extruded onto this structure to create 3D shells rather than the conventional ‘2.5D’ layers. Measured lengths of masking tape were cut and successively applied to the jig to act as a substrate. The adhesive area of the tape was minimised, to aid the removal of samples from the jig, by adhering sections of tape to one another.

The adhesive area of the tape was minimised, to aid the removal of samples from the jig, by adhering sections of tape to one another. Additional purposes of the tape were to: provide a rougher surface than the machined jig, and; provide thermal insulation between the molten polymer and the aluminium.The 5AXISMAKER was calibrated to ensure that measured linear and rotational movements corresponded with the readings of the CNC machine controller. Relationships between filament feed rate (controlled by the A axis stepper motor), NTC speed, and extruder temperature, were identified and these parameters were optimised via 3-axis ME tests. Stepper motor speeds that controlled X, Y, Z, B, and C axis speeds were finely tuned by executing the 5- axis G-code.

Stepper motor speeds that controlled X, Y, Z, B, and C axis speeds were finely tuned by executing the 5- axis G-code. Each 5-axis G-code required the completion of this process, as NTC speed is dependent on movement in all axes as well as the rate of curvature change in the tool path.To calibrate the X and Y axes, a locating pin was fastened to the machine tool collar and inserted into the corresponding hole on the top face of the jig (Fig. 7a). Once movement of the jig was constrained to the local Z axis, it was secured using a vice. This process aligned the machine Z axis with the jig Z axis, as well as zeroing X and Y coordinate values. Fine alignment was carried out by dry running the G-code with reduced motor speeds, after replacing the locating pin component with the extruder assembly.After the sample production setup and dry running processes were complete and motor speed settings were changed to the optimised parameter values, the generated 5-axis G-code was executed.

At this stage material was deposited onto the tape to form the first shell. All commands in the G-code were then followed to result in a 5-axis sample of a mechanically-optimised wave spring (as shown in Figure 7). The manufacture of each 5-axis sample required a complete repeat of this process, starting from tape application.2012abFigure 7 - A sample produced via 5-axis ME: (a) during manufacture, and (b) after manufacture3-Axis G-code GenerationMeasurements of all the 5-axis ME sample parameters (as shown in Figure 1) were taken to evaluate the differences between the manufactured geometries and the intended spring dimensions. These measurements informed the dimensions of the generated volumetric meshes that were used in the production of 3-axis ME samples.

These measurements informed the dimensions of the generated volumetric meshes that were used in the production of 3-axis ME samples. This process compensated for deviations between the 5-axis samples and their intended geometries.An additional functionality of the component cluster developed in Grasshopper 3D was its capability to generate closed, volumetric meshes that could be sliced using MakerBot Desktop software [ 17 ]. Two sinusoidal waves were generated to function as guide geometry for a rectangular cross section to be swept along, which formed a mesh of a single wave spring layer.

Two sinusoidal waves were generated to function as guide geometry for a rectangular cross section to be swept along, which formed a mesh of a single wave spring layer. Similarly to the 5-axis track generation method, copies of the mesh were rotated and translated according to the number of layers and number of peaks within the layer as selected using the developed component cluster (Figure 8a).abFigure 8 - Use of (a) the developed component cluster to generate (b) a closed mesh ready for G- code generation2013Meshes (Figure 8b) produced by the component cluster were then exported in STL format, and imported into a mesh editing software, Autodesk Netfabb Standard 2018 [18], to remove self-intersections and output a single, volumetric mesh. The modified mesh was then imported in MakerBot Desktop software [17] where slicing parameter values were selected, by attempting to match 5-axis and 3-axis parameter values where possible.

The modified mesh was then imported in MakerBot Desktop software [17] where slicing parameter values were selected, by attempting to match 5-axis and 3-axis parameter values where possible. The 3D mesh was processed to form a series of 2D areas, and the associated G-code was executed to manufacture 3-axis samples (a 3-axis ME example part is shown in Figure 9).Figure 9 - A 3-axis wave spring sample, including supporting structuresThe reasons for some parameter values not being identical were related to hardware variations, such as, the 5AXISMAKER and the MakerBot Replicator 2 extruder assemblies using 0.6 mm and 0.4 mm diameter nozzles, respectively. This difference resulted in a necessary dissimilarity between inter-track distances and inter-layer/shell thicknesses to maintain stable extrusion quality.

This difference resulted in a necessary dissimilarity between inter-track distances and inter-layer/shell thicknesses to maintain stable extrusion quality. Despite these differences, the measured weight was similar between sample sets.Table 1 records the average measured values of all fabricated samples, including: theparameters shown in Figure 1, the build time, and their masses.Table 1 - Build data for compression test samplesNumber of LayersSymbol Z Units 5-axis 3-axisN/A 4 4Number of Waves per Layer N N/A 4 4Thickness Meant mm 2.5 2.5Diameter Dm mm 37.4 37.3Radial Width b mm 7.7 7.6Build Time N/A min 97 110Compression TestingAxial compression tests were carried out using an Instron 3366 materials testing system with a 5 kN load cell and a compression speed of 2 mm.s-1. Load and deflection were measured until failure.

Load and deflection were measured until failure. The compression test was conducted to demonstrate the difference in stiffness and maximum compressive load between 5-Axis and 3-Axis wave spring samples. The experimental setup can be seen in Figure 10.2014MassN/A g 10.6 10.7Figure 10 - Compression testing setupResults and DiscussionThe compression test results from the three, 3-Axis samples produced using the MakerBot Replicator 2 machine and three, 5-axis samples produced using the 5AXISMAKER machine were exported to an Excel spreadsheet and used to plot graphs of load against deflection (shown in Figure 11). Results from only three samples were available due to time constraints. In both the 3-axis and 5-axis conditions, it can be seen that two of the samples performed similarly to one another while a third sample was significantly different. These two significantly different samples were treated as ‘outliers’.

These two significantly different samples were treated as ‘outliers’. For further analysis of the results and comparisons between the conditions, average values calculated only from the two similar pairs of samples were used.12001000)800N( d a o L60040020000246810Deflection (mm)Figure 11 - A graph of force profiles generated using compression testing dataIn terms of initial stiffness during small deflections (up to 1 mm), the 3-axis and 5-axis samples had very similar values of 240.3 N.mm-1 and 248.0 N.mm-1, respectively. This was to be expected since they were made from the same material and, under low loads, the impact of any layer-based differences in geometry or material properties would be negligible. However, above20155A-S15A-S25A-S33A-S13A-S23A-S3a load value of 300 N, the load-deflection curves for the 3-axis and 5-axis samples begin to diverge and continue to do so increasingly.

At a load value 600 N, the average deflection of the 3-axis samples is 0.45 mm greater than that for the 5-axis samples, i.e. they are exhibiting a lower stiffness. The precise reason for this reduction in stiffness is not yet understood but the authors suggest that it could be due the reduction in mechanical properties caused by incomplete material bonding between layers, as identified by Ahn et al (2002) [9]. In the 3-axis samples, the compressive load is transferred across many layer interfaces, whereas in the 5-axis samples, the load is largely contained within the curved shape of the layers. Hence, the anisotropic nature of the mechanical properties would have much more influence on the 3-axis samples.The second notable difference in performance between the two sets of samples is the higher ultimate strength and increased elongation to failure exhibited by the 5-axis samples. The average maximum load values for the 3-axis and 5-axis samples were 660 N and 975 N, respectively.

The average maximum load values for the 3-axis and 5-axis samples were 660 N and 975 N, respectively. This can be interpreted as the 5-axis samples being 47% stronger than the 3-axis samples. Also, the average elongation to failure was 3.6 mm for the 3-axis samples and 7.3 mm for the 5-axis samples, i.e. more than a doubling in the value. However, it should be noted that there was significant variation in the elongation to failure in the 5-axis samples. Never-the-less, both the 5-axis samples performed significantly better than the 3-axis samples. These data show that aligning the material extrusion direction to the mechanical stress vectors of the wave springs has resulted in a stronger and more resilient manufacturing solution.The third notable difference between the two sets of samples is not so apparent when looking at the recorded data, but was observed directly during the compression testing of the samples.

This difference relates to the nature of the failures in the 3-axis and 5-axis samples. The 3-axis samples all failed in a brittle, catastrophic manner, with the samples seeming to ‘explode’ in the test apparatus. In contrast, the 5-axis samples failed in a more prolonged, stepwise manner, with the samples remaining largely intact after the tests. Examination of the samples after testing revealed that the 3-axis samples had typically failed by several cracks being propagated through all the layers at more than one region of the wave spring (see Figure 12a). Similar examination of the 5-axis samples revealed that it was more typical for a crack to have propagated through some but not all layers in a smaller number of regions (see Figure 12b).

Similar examination of the 5-axis samples revealed that it was more typical for a crack to have propagated through some but not all layers in a smaller number of regions (see Figure 12b). The reason for this apparent increase in plasticity from the 3-axis to the 5-axis samples is not fully understood and will be the subject of further investigations.CrackFracture surfacesabFigure 12 - A comparison of wave springs produced via (a) 3-axis and (b) 5-axis ME2016As well as the samples fabricated via 5-axis ME exhibited superior mechanical properties relative to their 3-axis equivalents, the authors deemed them to be more aesthetically pleasing due the lack of the stair-stepping effect.

This advantage alone is unlikely to be sufficient reason for the use of 5-axis ME, due to the complexity of the process, predominantly: G-code generation and machine and jig setup processes.Although the design of the developed component cluster (Figure 3) allows for the most time-consuming G-code generation stage, the creation of the tool path, to be completed automatically, the cluster is limited to only producing a continuous track that describes a wave spring. One of the main appeals of 3-axis ME is the relative simplicity of the G-code generation process [19], as almost any volumetric mesh can be processed entirely automatically. For 5-axis ME to become a more viable option of manufacturing parts, with optimised mechanical properties, the tool path creation process should be more automated.

For 5-axis ME to become a more viable option of manufacturing parts, with optimised mechanical properties, the tool path creation process should be more automated. 3-axis ME strategies and machines that provide options of optimising mechanical properties of parts typically rely on increasing feature thicknesses or reinforcing layers with heterogeneous materials [ 20 ], respectively. To minimise the differences between the conventional G-code generation process for ME, the authors propose that volumetric meshes are similarly imported into a virtual environment, and at that point the user, or an additional automated process, determines the ME direction. This process may also incorporate jig production processes if required.

This process may also incorporate jig production processes if required. To compare the above described software approach with the process followed to produce a single wave spring highlights the gaps that need to be addressed for the adoption of 5-axis ME.ConclusionsThis work has demonstrated that 5-axis ME can be used as an effective means to generate high quality polymer parts. The samples created by the 5-axis process had improved surface finish, higher strength and a non-catastrophic failure mechanism. The results from the compression testing showed a 47% increase in strength, comparable to the 40% increase seen by Sinanemi et al in their experiments with curved layer ME [7]. This was due largely to the material extrusion tracks being aligned with the stress tensors experienced by the wave spring samples, a particularly novel aspect of this work. Deflection to failure was also increased significantly.

Deflection to failure was also increased significantly. The authors recognise that these results are tentative in nature since the sample size was small and a there was marked variability in the results. Never-the-less, they do indicate that 5-axis ME is a promising direction to pursue.The ‘downside’ of 5-axis ME is the added complexity of the equipment (including build set-up) and the additional time and effort needed to generate tool paths compared to those required for 3-axis ME. For the samples created for this paper, numerous steps were involved using several different software packages. This would be an unrealistic process chain for design and manufacturing engineers to follow in a commercial environment.At time of writing, further samples were being built for the purpose of confirming the initial compression test results.

Finite element analysis (FEA) will then be used to gain a better understanding of the failure modes and the reason for the apparent increase in plasticity in the 5- axis samples. Looking further ahead, a new project has been started with the aim of developing2017an industry-ready 5-axis ME system and associated software to facilitate automated generation of tool paths from design and FEA data [21].AcknowledgementsThis work was partly funded through the Engineering and Physical Sciences Research Council (Additive Manufacturing and 3D Printing Centre for Doctoral Training) and the European Commission (AMa-TUC Project 691787). Thanks also go to 5AXISWORKS Ltd for their help in programming and calibrating the 5-axis manufacturing system. Additional thanks go to the technical staff at Loughborough Design School for providing key equipment that enabled timely sample production.References[1][2][3][4][5]N. Turner, B., Strong, R., & A. Gold, S. (2014).

PHA_I, REMOTE MANUFACTURING, ANDSOLID FREEFORM FABRICATION *Samuel H. Drake University of UtahMay 1990work was supported in part by DARPA (N00014-88-K-0689). All opinions, findings, conclusions or recommendations expressed in this document are those of the author and do not necessarily reflect the views of the sponsoring agencies.114ALPHA_I, REMOTE MANUFACTURING, ANDSOLID FREEFORM FABRICATION *Samuel H. Drake University of UtahMay 1990AbstractAlpha_l is a nonuniform rational B-spline (NURBs) based solid modeling system that has been developed at the University of Utah over the past 10 years. In addition to being useful in modeling objects that are described by simple rotation and extrusion operations, the real power of Alpha_l is demonstrated in the modeling of complex parts with sculptured surfaces.

For the past several years, a major research thrust has been to use Alpha_l to semi-automatically generate process plan information and numerical control code to manu facture mechanical parts directly from the models. A long term goal is to support an on-line remote manufacturing facility for producing prototype parts. Recently, a 3D Systems stereo lithography machine has been added to the advanced manufacturing laboratory. The stereo lithography process and other SFF techniques are of particular interest for supporting a remote manufacturing facility in that these processes are inherently much safer than nu merically controlled machining. Special Alpha_l interfaces including a new slicing algorithm are being developed for the SFF machine use.

Special Alpha_l interfaces including a new slicing algorithm are being developed for the SFF machine use. By generating a SFF part directly from its NURBs description, Alpha_l should facilitate the manufacture of complex parts while providing smoother surfaces.1IntroductionAlpha_1 is a nonuniform rational B-spline (NURBs) based solid modeling system that has been developed at the University of Utah over the past 10 years. While Alpha_1 is quite useful in modeling objects that are described by simple rotation and extrusion operations, the real power of Alpha_1 is in modeling complex parts that include sculptured surfaces.For the past several years, a major research thrust has been to use Alpha_1 to semi automatically generate process plan information and numerical control code to manufacture mechanical parts directly from the models. A long term goal is to support an on-line remote manufacturing facility for producing prototype parts.

A long term goal is to support an on-line remote manufacturing facility for producing prototype parts. Ultimately, it should be possible to design"'This work was supported in part by DARPA (N00014-88-K-0689). All opinions, findings, conclusions or recommendations expressed in this document are those of the author and do not necessarily reflect the views of the sponsoring agencies.115parts or complete assemblies of parts in one location and have them manufactured in one or more remote locations and then shipped to where they are needed. This system would be a mechanical analog to the MOSIS (Metal Oxide on Silicon Implementation Service) that is now available. for vtsI.designers.Currently, the University of Utah advanced manufacturing laboratory has several state-of the-art CNC machines including a 5-.axis vertical machining center (milling machine) and a turn ing center (lathe) with a servo spindle and powered rotary tools.

A direct computer controlled coordinate measuring machine is being acquired to both automatically inspect manufactured parts and to input data from existing parts and physical models.Recently, a 3D Systems stereo lithography machine has been added to the advanced manu facturing laboratory. The stereo lithography process and other solid freeform fabrication (SFF) techniques have several major advantages for initial experiments in supporting a remote manu facturing facility. It is an inherently safe process in that errors can not cause a crash that could damage the machine or even injure the operator. Also, considerably less information is required to manufacture a part using SFF. With conventional machining, it is necessary to have com plete process planning information including information on sequence of operations, fixturing, cutting tools, and material properties.

With conventional machining, it is necessary to have com plete process planning information including information on sequence of operations, fixturing, cutting tools, and material properties. Conversely, with SFF there is no material information, no cutting tool information, and almost no fixturing information required.In many cases, the SFF process should allow prototype parts to be made without machining. This will be especially true if parts can be made from engineering materials either directly or indirectly. Even if more conventional manufactuing techniques are ultimately required, SFF processes can be used to prove out the model before using the conventional techniques.2 The Alpha_l Modeling SystemAlpha_l is a spline based, boundary representation solid modeler that is used to model a large class of sculptured and non-sculptured mechanical parts.

Objects which fall into this category include turbine blades, airplanes, helicopters, automobiles and automobile parts, sewing machine parts, robots and robot parts, molded plastic parts and the molds to produce them, etc. Even parts that have, at first glance, a simple geometry sometimes cannot be accurately modeled without sculptured surfaces.AlphaJ also demonstrates the utility of high quality graphics in the mechanical design process, particularly when dealing with sculptured surfaces. Proper perception of a complex shape is not possible from a line drawing or rough shaded rendering. Much of the perceived information about the shape of a sculptured object comes from the highlights caused by specular reflections from the surface. Without an accurate lighting model and a high resolution image, the highlights, and thus the perceived shape, will be distorted.

Without an accurate lighting model and a high resolution image, the highlights, and thus the perceived shape, will be distorted. Surface ripples that might be invisible in a line drawing show up immediately in a shaded raster image.AlphaJ also serves as an experimental testbed on which new modeling ideas and techniquescan be easily mounted and evaluated.1162.1 B·Splines in Alpha_1The tensor product parametric NURBs surface representation is the single representation that occurs throughout the Alpha.l system, and is supported by the entire system. Much of the system has been extended to support the "trimmed" surfaces, which are the results of Boolean operations derived. from the eSG design paradigm.B-splines have a nu.mber of desirable properties for use in computer-aided geometric design. A B-spline curve/surface is a piecewise rational function and exhibits specified degrees of con tinuity along specified isoparametric values, called "knot lines".

A B-spline curve/surface is a piecewise rational function and exhibits specified degrees of con tinuity along specified isoparametric values, called "knot lines". For example, cubic B-spline curves can be C 3 , C 2 , C 1 , Co, or discontinuous at such a knot value, as specified. This ensures that a curve which should be smooth is indeed smooth, but by properly manipulating the knot vector, discontinuities of specified derivatives may be inserted in the curve.B-splines have the convex hull property; that is, the curve always lies within the convex hull of the control points. It is therefore easy to bound the curve/surface, simplifying the calculations such as those required for producing an image or calculating errors in NC paths. Another feature of B-splines that is useful in a design system is that they provide "local control". Changing the location of a single control point modifies the shape of only a portion of the surface; the rest of the surface remains unchanged.

Changing the location of a single control point modifies the shape of only a portion of the surface; the rest of the surface remains unchanged. This allows, for example, a designer to shape one region of a surface without affecting other regions that may be already complete.The use of rational B-splines allows the exact representation of a large class of shapes, including the common quadric surfaces such as spheres, cylinders and cones. A polynomial B-spline, since it is a piecewise parametric polynomial curve or surface, can only approximate a circular or elliptical shape. By adding a weight or homogeneous component at each control point, it is possible to model these shapes exactly. Since many common mechanical parts have circular cross sections, it is critical that a design system be able to reproduce such shapes exactly. A B-spline may be evaluated by computing the value at any given point. Often, however, an approximation to the entire curve or surface is desired.

A B-spline may be evaluated by computing the value at any given point. Often, however, an approximation to the entire curve or surface is desired. This may be found via a "subdivision" or "refinement" process. The Oslo Algorithm provides a method for taking a B-spline and computing a new B-spline, point for point identical to the original, but with more knots, and therefore more control points and more polynomial spans. It can be used to split, or subdivide, a B-spline into two pieces that, together, make tip the original. The new control polygon, formed by joining the control points in order, more closely approximates the curve than did the original. Thus, by adding enough new knots, the control polygon itself may be used as an approximation to the curve.2.2 A Short Description of the Alpha_l SystemThe Alpha.J.. system provides a number of capabilities.

system provides a number of capabilities. A short listing of the features of Alpha_l includes:an interactive, extensible geometric editor,drafting style geometric construction operators,high-level shape modification and design operators,117automatic creation of primitive objects,support. for parametric design,feature based" design,the ability to perform set operations to combine objects into more complex assemblies,high quality graphical renderings of object models,computation of simple mass properties of modeled objects,direct interface to 3D systems stereo lithography machine, andmodel driven CNC code generation2.3 Set OperationsThe ability to perform set operations, usually union, intersection, and difference, on models is recognized as one of the characteristics of a true solid modeling system. Set operations perform a number of useful functions in the modeling process.

Set operations perform a number of useful functions in the modeling process. They can be used to combine simpler It is not necessary to design an entire object in one piece; objects into more complex ones. it may be broken down into simpler components that can be designed independently. The regions where the different parts meet are not explicitly designed, but arise from the interaction between the independently designed pieces. Using set operations to combine the parts produces the interface region automatically. A good example of this interaction occurs in a turbine blade, where the airfoil portion of the blade must be designed to meet certain aerodynamic or hydrodynamic constraints. The root of the blade is designed to hold the blade into the turbine during operation, and must meet an entirely different set of mechanical constraints.

The root of the blade is designed to hold the blade into the turbine during operation, and must meet an entirely different set of mechanical constraints. The airfoil must, however, be attached to the root, and a set operation is an ideal way to model the merger of the two.The Alpha_l system has the ability to perform set operations on "partially bounded sets". As long as they satisfy certain loose criteria, boundaries of objects to be combined by set operations need not be closed. This is a nice feature for the designer, who is not arbitrarily forced to close object boundaries just so that set operations may be performed.

This is a nice feature for the designer, who is not arbitrarily forced to close object boundaries just so that set operations may be performed. A good example of this is seen in the design of a turbine blade root, where none of the boundaries of the subpieces completely enclose a volume, but the final model does have a closed boundary.2.4 Graphicsin A key concept put forth by the Alpha_l project is that design of mechanical parts and, particular, of sculptured mechanical parts, is greatly aided by high quality graphic rendering of the model. Towards this end, the AlphaJ system incorporates both a scanline rendering program and a ray tracing program that can produce images of B-spline surfaces. This process can adaptively subdivide surfaces until the pieces are flat, to a given resolution, produce polyhedral approximations, and either ray trace or scanline render the resulting polygons.

This process can adaptively subdivide surfaces until the pieces are flat, to a given resolution, produce polyhedral approximations, and either ray trace or scanline render the resulting polygons. The progranl has many options to control the rendering process, the lighting model, surface characteristics,118and output form. It is possible to specify transparent or semi-transparent surfaces, allowing the designer to see interior details as well. Control over lighting is also provided, to allow careful placement of highlights to give maximum shape information. Images created with the ray tracing program may also include shadows, reflections, and texture mapping.Geometric EditorThe geometric editors for Alpha_I, shape_edit and model-server, are implemented in Portable Standard Lisp (PSL). Lisp provides a rich interpretive environment in which arbitrary and extensible data structures may be easily created.

Lisp provides a rich interpretive environment in which arbitrary and extensible data structures may be easily created. The embedding of the editor in a language leads easily to the concept of the program as the model, rather than the output created by the program. It is also easy, therefore, to create a parametric model, one that can represent a family of parts by changing a set of parameters. Many tools and operations are provided for shape definition. These include drafting operations, sweep operations, shape modification operations . and primitive objects.The drafting paradigm of design is supported by providing a number of drafting type opera tions that deal with points, lines, and arcs in the plane. Arcs and line segments may be chained together into a B-spline curve. Curves can be used to construct surfaces in many ways.

Arcs and line segments may be chained together into a B-spline curve. Curves can be used to construct surfaces in many ways. Alpha_l also has a particularly rich set of sweep operators that can be used to generate sculpted surfaces in a predictable manner.To simplify the task of working with B-spline surfaces, high level shape modification and design operators have been created. For example, the bend operator takes a previously defined surface and bends it. A thickening operation starts with a single surface, computes a surface that is offset from it by a given amount, and joins the edges of the two to form an object shaped like the original surface.To accommodate models created by CSG systems, or for those cases when primitive shapes are desired, routines are included for automatically creating some basic objects. These include boxes, spheres and ellipsoids, cylinders and cones, and tori.

These include boxes, spheres and ellipsoids, cylinders and cones, and tori. The surfaces of the primitives are represented as B-splines, so any of the shape modification operators can be applied to them. Real objects often have rounded corners, instead of the sharp corners on the primitive shapes. Therefore, provision has been made to create versions of some of the primitives with rounded corners. In particular, cuboidal shapes with rounded edges and corners are supported. The radii need not be identical on all edges, and the faces need not be at right angles.Finally, special purpose routines can be written to model any desired family of shapes. An example of this are the manufacturing features which can be optionally loaded into the geometric editor. Features including counterbores, countersinks, slotted holes, and rounded pockets have been defined in terms of built-in objects in the geometric editor.

Features including counterbores, countersinks, slotted holes, and rounded pockets have been defined in terms of built-in objects in the geometric editor. Linear and radial patterns of those features can also be constructed.3 Using Alpha_l for Solid Freeform FabricationSFF machines will be accessible to a potentially much larger user community than milling and turning machines. Therefore it is important to be able to drive them as automatically and as directly from the model as possible.119Special Alpha_l interfaces are being developed to drive the 3D SFF directly from the model. To implement this interface, it is necessary to generate a complete solid polyhedron whose facets are properly triangulated. This means that the process .requires triangles arranged so that an edge of one triangle cannotcontainave:rt~xofan.adjacent triangle.Fortunately, the scanline rendering algorithm produces polygonal approximations for both trimmed and regular NURBs surfaces.

Using that form, it is relatively easy to generate properly triangulated surfaces for models that have only a single surface or a collection of surfaces with correctly declared adjacencies. As most real models contain many surfaces, it becomes relative difficult for the designer to design surfaces· that exactly meet the other adjacent surfaces and provide the correct adjacency information. For complex parts it is generally easier to design simple overlapping surfaces and trim them using boolean or set operations. It should be noted Systems that it is sometimes possible to generate the effect of combining surfaces by using the " merge" operator.A more complex algorithm has been developed to properly triangulate models with the trimmed surface boundaries that result from using the boolean or set operators. Several recent parts represent the output from this algorithm.

Several recent parts represent the output from this algorithm. One potential problem with the current method of treating trimmed surfaces is that the current slicing .routine used by 3D Systems tends to collapse long thin triangles into degenerate lines.Anticipating potential future interfaces for SFF machines, we have also investigated the requirements and algorithms that would provide the necessary slice and layer information di rectly from the NURBs model and have demonstrated the ability to directly generate oriented slice curves from complex AlphaJ NURBs models. By generating a SFF part directly from its Alpha_l NURBs description, the manufacture of complex parts would be facilitated while providing more accurate surfaces.ReferencesDrake, S. and S. Sela, A Foundation for Features, Mechanical Engineering, Volume 111, No.1, January 1989.Riesenfeld, R.F., E. Cohen, and S. H.

Therefore, creating an en-hanced understanding of thermal phenomena through measurements and simulations is crucial to in-crease the reliability of the technology. Current research is mainly based on temperature measurementsof the upper layer, leaving major scope for the conditions at the substrate-part-interface. This area is ofutmost technical importance because it serves as the main heat sink. Insufﬁcient heat dissipation leadsto accumulations of heat, deformations, and process breakdowns. This contribution presents a simpleand ﬂexible method to analyze the thermal progression close to the part inside the substrate plate. Theacquired data shows very high consistency.

This contribution presents a simpleand ﬂexible method to analyze the thermal progression close to the part inside the substrate plate. Theacquired data shows very high consistency. Additionally, the results are compared to a model createdusing an ISEMP developed FEM-Software which shows promising results for validation studies.Keywords: Laser Powder Bed Fusion of Metals (LPBF-M), Thermocouple, Finite element modeling(FEM), Validation, Thermocouple, Stainless steel 1.4542 (GP1)1. IntroductionThe laser powder bed fusion of metals (LPBF-M) process has gained a lot of attention in research andindustry during the last decade and creates an increasing number of industrial business cases. It enablesbuilding-up complex metallic parts in a one-step process.

It enablesbuilding-up complex metallic parts in a one-step process. The potential for individual, geometricallyfunctional and optimized parts through the layerwise build up principle has been proven in multiplestudies [1–3] and led to immense growth in this business sector [4].One of the primary advantages of the LPBF-M process is the possibility to use innovative new materialswith tailored properties as well as build complex and demanding structures. However, this poses oneof the biggest challenges for this technology. Many promising materials, like maraging steel (MS1),high entropy alloys, and amorphous metals, are very sensitive towards their thermal history [5, 6].

A simple redesign of the substrate plate was made to enablemeasurements with only 2 mm distance to the part. Hence, it reaches a very high resolution while largelykeeping the integrity of the thermal system of the substrate plate.2. Experimental SystemThe experiments were conducted using an EOS M290 LPBF-M machine with modiﬁed substrate andbase plates as well as an integrated additional temperature sensor system. A regular substrate plate hasa surface area of 250 x 250 mm. Here, a C45 steel substrate plate of 220 mm x 220 mm was ﬁxatedon an adapter ring with minimal contact area. The base plate used has a recess, so that wiring for thethermocouples can be led from below the substrate plate through the shaft of the building platform.N30 [mm]Figure 1: Schematic close up of the additive part on the disc positioned in the slot within the substrate plate. The sensors S1 and S2, the diameter d, and the height h of the part are depicted.

The spatial domain Ω(t) ⊂ R3, which describes the already exposed solid material as well as the substrate plate, evolves in time due to the material addition in LPBF-M processes. It is assumed that Ω(t) is constant on each interval (ti,ti+1) and the addition (cid:26) of material happens instantaneously at time ti, i.e. Ω(t) = Ωi := k=0 Pk ∪ S for t ∈ [ti,ti+1), whereby i S describes the geometry of the substrate plate and Pk of the part layer k. The simulation model is based on the well known heat equation, which arises from combining energy balance with Fourier’s law.

Therefore, (1)is discretised via an implicit Euler scheme in time and the Finite Element Method (FEM) in space (formore information about solving parabolic equations by FEM see e.g. [22] from a mathematical or [23] from an engineering point of view).

The value was ﬁtted, so that the measured peak temperatures of the ﬁrst layer at S1 of the cylinder sample (d= 10 mm) are reproduced correctly. The measured peak temperatures show a highscattering due to a low time resolution (cf. section 4). Therefore, the peak temperature of the ﬁrst layer was estimated to be 83 ◦C by manually estimating a curve, which connects the measured maximal peak temperatures. The obtained activation energy density is 25.35 J/mm3, which is used for all samples. Note that the whole energy is inserted in one single layer. It is well known that several underlying layerswill be molten during the exposure of a new single layer in LPBF-M processes (this is important to guar-antee the connectivity of each new layer to the part and reduce the porosity).

Since the energy, which isnecessary to melt several layers, in reality, is inserted in the simulation into a single layer by activating it at θact, it is clear that the resulting temperatures θact are signiﬁcantly higher than melt temperature. The material parameters were taken from data sheets (substrate plate material: C45 steel [25], part material:GP1 / 1.4542 steel [21]). Only the speciﬁc heat capacity of GP1 / 1.4542 steel was generated with thehelp of the JMatPro software since it was not provided. The temperature dependent function of ther-mal conductivity is deﬁned as a linear inter-/extrapolation of the stated values.

The temperature dependent function of ther-mal conductivity is deﬁned as a linear inter-/extrapolation of the stated values. All used parameters aredescribed in table 1.λpart λsub-plate cρ,part cρ,sub-plate ρpart ρsub-plate[W/(m K)] [W/(m K)] [J/(kg K)] [J/(kg K)] (cid:6) (cid:5) kg/m3 kg/m3(cid:5)(cid:6)I20 ◦C 100 ◦C 200 ◦C 300 ◦C II II13141516II 4250047078007800Table 1: Material parameters used for the simulations1271The duration to recoat a new powder layer was reported by the used machine to be 9 s. In case of thecylindrical samples the exposure time of each layer was determined by measuring the overall processtime. After dividing it by the number of process layers and subtracting the recoating time one obtains the exposure time. This results in 1.2 s for the cylinder with d = 10 mm diameter and 0.6 s for the cylinder with d = 5 mm diameter.

This results in 1.2 s for the cylinder with d = 10 mm diameter and 0.6 s for the cylinder with d = 5 mm diameter. Since ﬁrst and last layer of the cone samples are equal to one of the cylindricalspecimens, the layer exposure time is assumed to be a linear in-/decreasing function of these times incase of the cone samples. Exemplary timing measurements agreed with this assumption.4. ResultsSix different experiments have been conducted to analyze four potential inﬂuencing factors on the tem-perature progression at the connection of the part to the substrate plate: Part volume, interface area tothe substrate plate, the inﬂuence of the copper paste, layer height.

This effect is more signiﬁcant forthe maximum peaks because of comparably low laser exposure time in contrast to recoating time. Therapid cooling process after the build up was also recorded. It can be used for ﬁtting the heat loss of thetotal system in the simulation. However, this approach is not further investigated for this work.BO75- 5/5 SI ----- 515 S27065~ 60 ~ ass e ~ so C. E ~ 45403530:: ,. " = ., ,. 3 ,. =25757065~ 60~1-- 10/10 Simil - 5/5Siml003 = " =Q~ 55 e ~ so C. E ~ 454035302500:0000:1500:3000:45 Time lhl01:0001:1500:0000:1500:3000:45 Time lhl01:0001:1501 :30Figure 6: Temperature progression of sensor 1 (S1) 2 mm below the part and sensor 2 (S2) under the substrate plate for the build up of a 10 mm high cylinder with a diameter of 10 mm (top left) and 5 mm (top right).

Reasons for the differences between simulation and measurement will be discussed insection 5.The second comparison (Figure: 7) shows the inﬂuence of energy input (∝ surface area) per exposure on the temperature progression. Therefore, two cones with same volume but reverse progression of diameter (d1 = 10 −→ 5 mm /d 2 = 5 mm −→ 10 mm) were build. Both temperature progressions start analogue to their cylindrical counterpart (Compare: 6(b/c)). The deviation from this similarity can be seen by a decrease of the average temperature (44 ◦C −→ 31.6 ◦C) for the ﬁrst and an increase (31.6 ◦C −→ 44 ◦C) for the second sample. Hence, the results show a clear correlation between the average ﬁnal temperatureand the area of exposure.The cone samples have also been simulated and the calculated temperatures at the position of S1 areshown.

From a qualitative point of view, a good agreement of measured and simulated temperatureprogression can be observed, as in the case of cylinder geometries. In particular, differences causedby changes in geometry/orientation are reproduced.

This effect is signiﬁcantly ampliﬁed for the sample without copper paste so that the increase in temperature is much faster, reaching 92.2 ◦C before decreasing. On the other hand, the minimum quite accurately resolves the same progression aswith copper paste. This leads to the conclusion, that the copper paste decreases the time needed to depletethe energy from a single laser exposure, but the recoating time is still long enough so that the effect isnot accumulating.

50 e ~ 4540' , · ...... ·-.·..:: ___ _ ___ J / 10 ___ __ ____ .. · - .10/54010/10355/5 - - - - ~ ~-~·.:.-.:· ::.:.·.; ~3530302500:0000:1500:3000:45 Time ihl01:0001 :1501:3000:0000:1500:3000:45 Time lhl01:0001:1501 :30Figure 10: Fitted progression of the maximum peak temperatures of sensor 1 (S1) for cylinder with d = 10 mm & d = 5 mm and cones with d = 10 − 5 mm & d = 5 − 10 mm for measurements left and simulation right.5. DiscussionThe experiments show very insightful results. Nevertheless, a few optimizations of the setup should bemade for further investigations.The ﬁxation of the thermocouples is a time-consuming procedure. The point weld must have good con-tact with the disc to ensure a decent thermal connection but also ﬁxate the stiff wire.

The point weld must have good con-tact with the disc to ensure a decent thermal connection but also ﬁxate the stiff wire. Since the results show a maximum temperature below 100 ◦C a heat resistant reversible glue could be used additionally to decrease setup time between experiments.The third comparison showed a signiﬁcant inﬂuence of copper paste on the overheating effect of thedisc in the ﬁrst few minutes. The thermal resistance between disc and plate still poses a deviation fromthe regular process and should be reduced. The simulations treat disc and substrate plate as one ho-mogeneous material. Hence, the disc heats up signiﬁcantly more for the experiment than predicted bysimulation. Furthermore, the simulated peak temperatures of each layer are decreasing faster than themeasured ones. This can be explained by the fact that the heat conduction from the disc downwards intothe substrate-plate is slower than in solid material.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceAluminum Matrix Syntactic Foam Fabricated with Additive ManufacturingM. Spratt1-a, J. Newkirk1-b, K. Chandrashekhara1-c1Missouri University of Science and Technology, 300 W. 14th St, Rolla, MO 65409amsf998@mst.edu, bjnewkirk@mst.edu, cchandra@mst.eduKeywordsAluminum, Lightweight, Composite, DEDAbstractSyntactic foams are lightweight structural composites with hollow reinforcing particles embedded in a soft matrix. These materials have applications in transportation, packaging, and armor due to properties such as relatively high specific stiffness, acoustic dampening, and impact absorption. Aluminum matrices are the most widely studied of metal matrix syntactic foams, but there is little to no research in regards to processing the foams with additive manufacturing.

Aluminum matrices are the most widely studied of metal matrix syntactic foams, but there is little to no research in regards to processing the foams with additive manufacturing. It is theorized that the fast cooling rates and limited kinetic energy input of additive could reduce two issues commonly associated with processing syntactic foams: microsphere flotation in the melt and microsphere fracture during processing. In this study, 4047 aluminum blended with glass particles was deposited on a 4047 Al substrate using an additive process. Characterization of the foams include mechanical testing and microstructural analysis.IntroductionThe aluminum composite built in this study is an aluminum matrix syntactic foam. Metalmatrix syntactic foams (MMSF) are lightweight structural composites composed of a metal matrix and rigid, hollow spheres. The hollow spheres embedded in the matrix significantly reduce the weight of the composite, while also offering some.

The hollow spheres embedded in the matrix significantly reduce the weight of the composite, while also offering some. The unique microstructure of the MMSF also offers other benefits such as vibrational dampening and impact absorption. When comparing the MMSF to a metal foam or a particulate composite, there is an improvement to the specific stiffness and specific strength. [1]Most of the literature surrounding syntactic foams focuses on refining two mainprocessing methods – stir casting and pressure infiltration. [1] Both methods include harsh mechanical forces that can easily fragment the delicate microspheres. Microsphere fragments can cause unwanted densification of the microstructure and possibly create stress concentrations in the matrix which would decrease the mechanical properties of the composite. It is possible that AM could form the material without applying significant mechanical forces to the microspheres.

It is possible that AM could form the material without applying significant mechanical forces to the microspheres. Microsphere buoyancy in the melt can also be countered by laser AM as the fast cooling rates and layer-by-layer nature should lock microspheres into layers instead of allowing them to float to the top of the part. Finally, AM is a near net shape fabrication process, a requirement for syntactic foams.242Aluminum alloys are difficult to laser additive manufacture for several reasons. Aluminum tends to be difficult to impossible to weld due to the high thermal conductivity of the metal (about 41% IACS for 4047 aluminum [2]) and the high surface tension of the liquid aluminum. Aluminum alloys also tend to reflect, rather than absorb, laser energy. Because of this, although aluminum alloys can be made with laser AM [3] [4], alloys tailored to be less reflective and conductive (such as 4047) tend to be more successful.

Because of this, although aluminum alloys can be made with laser AM [3] [4], alloys tailored to be less reflective and conductive (such as 4047) tend to be more successful. The Missouri S&T LAMP group built 4047 aluminum on their custom-built Directed Energy Deposition (DED) system. This study builds on that work done by Isanaka et al. [4] to laser additive manufacture an aluminum composite on the same DED system.In this project, a syntactic foam with a 4047 aluminum alloy and hollow borosilicateglass microspheres was fabricated with the Missouri S&T custom-built DED system. This was done by first finding the optimal DED parameters for 4047 aluminum with solid SLS glass particles. Specimens were then optically observed for intact microspheres. This work is a continuation of work presented by the author at the MS&T conference in 2016. [5]Experimental ProcedureThe DED equipment used in this project was custom built by the Missouri S&T LAMP group students.

[5]Experimental ProcedureThe DED equipment used in this project was custom built by the Missouri S&T LAMP group students. This machine is still in the process of equipment optimization. Figure 1 shows a schematic of the equipment setup. The substrate used in this experiment was a 1”x0.5”x0.125” sheet of 4047 aluminum from Eagle Alloys Corp.Figure 1: Schematic of the Missouri S&T DED system built by students.4047 aluminum powder from Valimet Inc. was used in this experiment. The aluminum had a D50 of 70 microns according to the manufacturer. [6] Two types of glass were used. The first was a soda-lime-silicate (SLS) glass that was crushed into a 45 to 75 micron particle size range. This was done by ball milling large glass particles for 24 hours, then sieving to the correct size. The second type of glass was a borosilicate glass from MO-SCI LLC. This material consisted of hollow glass spheres between 1-140 microns in size.

The second type of glass was a borosilicate glass from MO-SCI LLC. This material consisted of hollow glass spheres between 1-140 microns in size. The glass was then sieved to the same 45 to 75 micron particles size range as before.243The first composition tested was 4047 aluminum with 30 vol.% SLS glass. This was attempted first in order to understand the laser parameters required to deposit the aluminum composite. The second composition was 4047 aluminum with 15% borosilicate hollow glass spheres. Table 1 shows the final parameters used tuning the machine for each composition. Table 2 shows the compositions of the 4047 aluminum, SLS glass, and borosilicate glass.

Optical microscope images and SEM imageswere taken. Point EDS data was collected in key areas.Results and DiscussionThe first section of this study was involved with the optimization of DED parameters fora material with 70% aluminum and 30% filler. This was done by depositing a composite comprised of 70% aluminum and 30% (by volume) SLS glass particles. The SLS glass was chosen because it was readily available, inexpensive, and nearly the same density as the aluminum at approximately 2.5 g/cc. The similar densities and particle size between the two materials reduced powder segregation. The SLS glass differed negatively from the borosilicate spheres in several important ways, however. First, the SLS glass was not perfectly spherical as it was made by ball milling. This caused the powder mixture to not flow well, hindering the244powder feeding process during deposition.

First, the SLS glass was not perfectly spherical as it was made by ball milling. This caused the powder mixture to not flow well, hindering the244powder feeding process during deposition. Second, this soft SLS glass was designed to melt at relatively low temperatures and has a wide temperature range under which is stays at a relatively low viscosity. [7] This resulted in the glass powder melting during laser sintering and flowing during processing. Finally, the SLS glass was not hollow.The laser parameters were tuned to this composition by building single track walls on thesurface of the substrate. A balance between the laser power and scan speed was found. A slow scan speed helped to keep a stable melt pool, while a low laser power helped to prevent overheating the melt and the substrate.Figures 1 and 2 show images of the aluminum and SLS glass composite microstructure. From these images, it is clear that the glass did melt.

From these images, it is clear that the glass did melt. EDS was also done on this sample in order to confirm the chemistries. The optical image (figure 2) shows that the structure of the aluminum is primarily a eutectic structure with aluminum precipitates. This is the case across all the specimens.SLS glassaluminumFigure 2: Unetched SEM image of the SLS glass and aluminum specimen. The phases are indicated on the figure.245poreSLS glassalpha aluminum precipitatealuminum eutectic structureFigure 3: Unetched optical image of the SLS glass and aluminum specimen. The phases are indicated on the figure.The second section of this study was the deposition of a powder mixture containing hollow borosilicate glass spheres and aluminum. Due to the very low density of the spheres (0.15 g/cc), the powder can easily segregate when agitated. Because of this, only 15 % borosilicate glass microspheres could be mixed into the aluminum without immediate segregation.

Because of this, only 15 % borosilicate glass microspheres could be mixed into the aluminum without immediate segregation. The parameters had to be further tuned to the change in filler content. This was done in the same manner as before.Figure 3 shows an optical image of one microsphere within the specimen. There werevery few microspheres within the specimens, and most of them were near the outer surfaces of the aluminum deposit. Serial polishing through 1 mm of the specimen, in intervals of approximate 50 microns, revealed less than 10 microspheres per specimen. It is likely that the lightweight microspheres simply do not have the kinetic energy to penetrate the aluminum melt pool, a necessary step of the DED process. It is also possible that some of the microspheres fully melted in the aluminum melt pool.

It is also possible that some of the microspheres fully melted in the aluminum melt pool. A powder bed process may be a better method for fabricating syntactic foams as the microspheres would not have to inject into the meltpool.246Borosilicate glassaluminum eutectic structureporeFigure 4: Optical Image of the borosilicate glass and aluminum specimen. Phases on indicated on the figure.Conclusions and Future WorkThe major conclusion to be drawn from this work is that additive manufacturing of an aluminum matrix syntactic foam is possible, but not without great difficulty. A powder bed method may work much better for this composition, to be certain that the microspheres enter the melt pool. Future work for this project includes the utilization of a powder bed, as well as increasing the volume fraction of the microspheres from 15 to 30 volume %.References[1] P. K. Rohatgi, N. Gupta, B. F. Schultz and D. D.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAN ALUMINUM-LITHIUM ALLOY PRODUCED BY LASER POWDER BED FUSIONYang Qi*, Hu Zhang†, Haihong Zhu*, Xiaojia Nie*, Xiaoyan Zeng* *Wuhan National Laboratory for Optoelectronics, Huazhong University of Science and Technology, Wuhan, Hubei 430074, PR China †School of Optical and Electronic Information, Huazhong University of Science and Technology, Wuhan, Hubei 430074, PR ChinaAbstractAluminum-lithium (Al-Li) alloys are promising to replace traditional high-strength aluminum alloys in aerospace and military industries due to their low density, high specific strength, and excellent corrosion resistance. However, there is little research focused on the laser powder bed fusion (LPBF) of Al-Li alloys due to their poor weldability and high crack susceptibility.

However, there is little research focused on the laser powder bed fusion (LPBF) of Al-Li alloys due to their poor weldability and high crack susceptibility. In this study, the feasibility of the Al-Li alloy fabricated by LPBF was investigated. The effect of the processing parameters on the densification and crack formation behavior was studied. Finally, after optimizing the processing parameters, crack-free and nearly fully dense 2195 Al-Li alloy 3D- printed samples were obtained. A relative density of 99.92% and a microhardness of 89.1 HV were achieved.IntroductionAluminum-lithium (Al-Li) alloys are of great interest to the aerospace and military industries because they offer many advantages over conventional aluminum alloys [1]. Their low density, high specific strength, and excellent corrosion resistance can bring significant weight savings [2], leading to Al-Li alloys being increasingly substituted for conventional 2xxx and 7xxx series high-strength aluminum alloys [3].

The 2195 Al-Cu-Li-Mg-Ag alloy, one of the representative latest generation of aluminum-lithium alloys, has been used in the production of the lightweight parts in aerospace [4], due to its excellent strength at room temperature [5] and high performance even at cryogenic atmosphere [6].including casting, pressing, and extrusion are complex, time-consuming and unable to produce complex structures [7]. The laser powder bed fusion (LPBF), also known as selective laser melting, is an additive manufacturing technology with design freedom and manufacturing flexibility. The LPBF has the advantages of short cycle times, low tooling costs, and eliminating consumption-of-scale constraints [8][9]. Nowadays, the printable materials of the LPBF are focused on the steel [10][11], nickel-based superalloy [12][13], titanium alloy [14][15], and aluminum-silicon alloy [16][17].

Nowadays, the printable materials of the LPBF are focused on the steel [10][11], nickel-based superalloy [12][13], titanium alloy [14][15], and aluminum-silicon alloy [16][17]. The 2xxx and 7xxx series high-strength aluminum alloys with poor weldability have also been reported to be successfully fabricated by the LPBF [18][19][20][21][22]. Little research is focused on the LPBF of Al-Li alloys till now. Al-Li alloys are susceptible to hot cracking during laser processing due to their high coefficient of thermal expansion and eutectics formation during solidification [23].However, the traditional manufacturing methods of Al-Li alloys,This work investigated the feasibility of the 2195 Al-Li alloy fabricated by LPBF.

The effect of the scanning velocity, hatch spacing, and the volumetric energy density on the densification and crack formation behavior was studied.609Materials and methodsGas-atomized 2195 Al-Li powders supplied by Intelligence Manufacturing Technology were used in this study. The chemical composition listed in Table 1 was analyzed by Inductively Coupled Plasma Optical Emission Spectrometry (ICP-OES). The powder morphology was observed by an FEI Nova Nano SEM 450 scanning electron microscope (SEM) as shown in Fig. 1a. The raw powder consists of many spherical powder particles and some small aggregated satellite particles attached to the big particles. The powder particle size distribution shown in Fig. 1b was measured by Malvern UK Mastersizer 3000.Table 1.

The powder particle size distribution shown in Fig. 1b was measured by Malvern UK Mastersizer 3000.Table 1. Chemical composition (wt.%) of the powder used in this studyElement PowderCu 3.95Li 1.08Mg 0.68Ag 0.40Zr 0.11Si 0.05Al Bal.(b) 1210010~ 8 ~ C 0 ti 6 ~ E 4 :::, 0 > 2Q)0D10 = 18.?µm D50 = 35.3µm D90 = 58.9µm80 ~ u 111 .::C 060Q) E :::, 40 0 > Q) >20 ~ :5 E :::, 00110100Particle size (µm)Fig. 1. (a) SEM image showing the morphology of the raw powder. (b) Powder particle size distribution.The LPBF experiment was carried out at room temperature on a self-developed LPBF machine (LSNF-I). The LPBF machine consists of a continuous wave IPG YLR-200 fiber laser with a maximum laser power of 200W, and other details have been given elsewhere [24][25]. All samples were fabricated under a protective argon atmosphere with the concentrations of H2O and O2 controlled below 100 ppm.

All samples were fabricated under a protective argon atmosphere with the concentrations of H2O and O2 controlled below 100 ppm. To investigate the feasibility of the Al-Li alloy fabricated by LPBF, 8 mm (length) × 8 mm (width) × 4 mm (height) cubic samples were built with different processing parameters. The laser power, layer thickness, and the layer-to-layer rotation angle were fixed at 200 W, 30 μm, and 90°, respectively. Other processing parameters are shown in Table 2.Table 2. LPBF processing parameters used in this studyParameters (Unit)ValueScanning velocity (mm/s) Hatch spacing (mm)80, 160, 240, and 320 0.06, 0.08, 0.10, and 0.12The relative density of every cubic sample was obtained by image processing of six optical microscope graphs using Image-Pro Plus 6.0 software. All samples were subjected to a standard metallographic procedure.610Results and discussionsFig.

All samples were subjected to a standard metallographic procedure.610Results and discussionsFig. 2 shows the effect of the scanning velocity on the relative density with a fixed hatch spacing of 0.12 mm. The relative density decreases as the scanning velocity increases. For every 80 mm/s increase in the scanning velocity, the relative density decreases more. The nearly fully dense 2195 Al-Li alloy sample is obtained when the scanning velocity is 80 mm/s.100.0'::f2.99.599.0~ ~ 98.5 'in C: (I) "'O (I) > ~ Q) ~98.097.597.096.596.0080160240320400Scanning velocity (mm/s)Fig. 2. Effect of the scanning velocity on the relative density.Fig. 3 presents the optical microscope graphs of polished samples fabricated with different scanning velocity at a fixed hatch spacing of 0.12 mm. The irregular pores are found to be in all the samples, and the size of the irregular pores increases with the increase of the scanning velocity.

6. Effect of the volumetric energy density on the relative density and the crack formationConsidering the porosity and the fabricating efficiency, the optimized parameters are the laser power of 200 W, the layer thickness of 30 μm, the scanning velocity of 80 mm/s, and the hatch spacing of 0.1 mm. The relative density and the microhardness of the sample built with the optimized parameters are 99.92% and 89.1 HV respectively.ConclusionsThe feasibility of the 2195 Al-Li alloy fabricated by LPBF has been successfully proved. The sample fabricated with the optimized parameters is crack-free and nearly fully dense with a relative density of 99.92% and a microhardness of 89.1 HV. The volumetric energy density is crucial for determining the crack formation.

The volumetric energy density is crucial for determining the crack formation. Crack-free samples can be obtained when the volumetric energy density is above 520 J/mm3.Future worksFuture works will focus on the tensile performance and the heat treatment of the 2195 Al-Li alloy fabricated by LPBF.AcknowledgmentsThis work was supported by the National Natural Science Foundation of China (Grant No. 51805184). The authors would also like to thank the Analytical and Testing Center of HUST for the SEM analysis.References[1] R.J. Rioja, J. Liu, The Evolution of Al-Li Base Products for Aerospace and Space[2]Applications, Metall. Mater. Trans. A. 43 (2012) 3325–3337. doi:10.1007/s11661-012- 1155-z. T. Dursun, C. Soutis, Recent developments in advanced aircraft aluminium alloys, Mater. Des. 56 (2014) 862–871. doi:10.1016/j.matdes.2013.12.002.613[3][4][5][6][7][8][9]B.I. Rodgers, P.B.

 Feature and Knowledge Based Process Planning for Additive Manufacturing in Multiple Parts Production ContextY. Zhang, A. BernardIRCCyN, Ecole Centrale de Nantes, 44300 Nantes, FranceAbstractAdditive Manufacturing (AM) has played an important role in manufacturing, especially in customized production. It is an ideal 'Concurrent Manufacturing' which enables fabricating a group of same or even different multiple parts simultaneously within one build volume due to its unique layer by layer processing way. However, there is very few available methods or tools for users, e.g. the AM manufacturing service bureaus, to optimize the process and production plan in multiple parts production context. To deal with this problem, this paper introduces an AM feature and knowledge based systematic process planning strategy. The main contents and key issues of process planning for AM in multiple parts production context are analyzed.

The main contents and key issues of process planning for AM in multiple parts production context are analyzed. Then, a developing CAPP system based on a systematic process planning framework for AM in this multiple parts production context is presented. Finally, some test examples are applied to demonstrate the functions and effectiveness of some key modules of the developing system.IntroductionAdditive Manufacturing (AM) processes employ a layer by layer processing manner to deposit material layers progressively for building 3D parts according to sliced 3D CAD models. This special processing way allows AM machines to build multiple parts simultaneously by placing multiple part slices within each building layer. Significant savings in cost and time can be achieved in rapid prototyping (RP/AM) by manufacturing multiple parts in a single setup to achieve efficient machine volume utilization [1].

Significant savings in cost and time can be achieved in rapid prototyping (RP/AM) by manufacturing multiple parts in a single setup to achieve efficient machine volume utilization [1]. Hence, in real application, to improve the machine utilization, parts are usually built batch by batch in AM machines per run but not only one by one. This forms the multiple parts production context (Figure 1).Figure 1, Difficulties of decision making in Multi-part production context.1259To make a group of parts built, process technicians or planners have to do a set of optimizations and decision makings for preparation work, process planning. As defined by Marsan and Dutta [2], there are usually four main planning tasks, orientation optimization, support design, slicing and tool-path/scanning-path planning. To solve these tasks, researchers had proposed numerous solutions [3-9].

To solve these tasks, researchers had proposed numerous solutions [3-9]. However, most of these solutions were designed for single part production context where only one part is built per machine run and they only focused on the operational level to help transfer a virtual CAD model to a physical model [3]. In addition, these solutions lack systematic. Systemic discussed here refers to two main aspects: the integrity of process planning content, including both of ‘macro planning’ (e.g. manufacturability analysis, etc) and ‘micro planning’ (specific planning tasks for processing, e.g. orientation optimization), the systematic analysis of the interdependence between different process planning tasks. Obviously, incomplete process planning cannot realize the full processing chain (Figure 2). The lack of interdependence analysis between the process planning tasks cannot guarantee optimal planning results. As stated by Kurkarni et al.

The lack of interdependence analysis between the process planning tasks cannot guarantee optimal planning results. As stated by Kurkarni et al. [3], ‘process planning problems are not individual problems alone, but they are related to each other’. Besides, in multiple parts production context, the planning tasks and their characteristics are different to those in single part production context. Hence, to realize a full processing chain and obtain better planning results for AM in multiple parts production context, this paper introduces a feature and knowledge based systematic process planning strategy.

To reduce the complexity, the post processing in the processing chain is not included.Figure 2, General processing chain of AM (‘micro level’; ‘one-way’ information flow).The left of this paper is arranged as follows: the second section will analyse the process planning problem in multiple parts production context; the third section will introduce the development of the proposed strategy; the fourth section will present the implementation of two key modules; the fifth section will present test examples for demonstration; the last section is the conclusion.Problem descriptionIn AM service bureaus (Figure 1), different orders may come from different clients with different production requirements, e.g. lead time, cost, quality, etc. Therefore, production technicians or planners have to make different decisions and optimizations so as to do the preparation work for production.

lead time, cost, quality, etc. Therefore, production technicians or planners have to make different decisions and optimizations so as to do the preparation work for production. The preparation work is mainly used to answer two types of key questions: ‘Whether a part is suitable to be processed by AM processes?’ and ‘How to produce a part?’ In this paper, all of the preparation work for AM production is defined as process planning and is proposed to be grouped into two levels: ‘macro planning’ and ‘micro planning’, which are used to deal with the two types of questions. In the ‘macro Planning’1260level, the main planning tasks include: manufacturability analysis, selection of AM process or manufacturing scenario, prediction of build time, cost and general part quality, etc. These tasks are usually beyond the processing chain in AM.

These tasks are usually beyond the processing chain in AM. While in the ‘micro planning’ level, the planning tasks are composed by: orientation optimization, work space planning, support generation, slicing, tool path planning, etc. To conduct process planning in multiple parts production context, the foremost thing is to identify the planning tasks. In AM service bureaus, different AM machines already install different preprocessing software tools to help process planners to deal with the planning tasks, but only for the ‘micro planning’ level. For the planning tasks in the ‘micro planning’ level, they usually use different algorithms designed for specific processes. It is hard to develop compatible support generation, slicing and path- planning algorithm for all the AM processes. Therefore, it is better to use the available tools to deal with the above three planning tasks.

Therefore, it is better to use the available tools to deal with the above three planning tasks. However, these tools cannot provide enough support to the orientation optimization and work space planning in the multiple parts production context. For example, current software (e.g. ‘Magics’ software) can only orientate parts one by one, which cannot guarantee an optimal orientation results. In multiple parts production context, the total build time, cost and quality not only depend on the individual part’s build orientations but also their combination. For space planning, current methods mainly use part’s bounding box and apply BL-GA methods [10], which waste much building space and cannot obtain optimal solutions. Fortunately, current AM machines usually can accept CAD models with STL format.

Fortunately, current AM machines usually can accept CAD models with STL format. This enables the availability of developing generic orientation optimization and work space planning methods with uniform output, positioned multiple CAD models in STL format, for the later planning tasks, e.g. support generation, slicing, on different preprocessing platforms installed in different AM machines. Therefore, there is no need to develop planning methods for all the planning tasks in the ‘micro planning’ level except for the orientation optimization and work space planning tasks. Hence, the general planning tasks in multiple parts production context can be identified as shown in Figure 3.

Hence, the general planning tasks in multiple parts production context can be identified as shown in Figure 3. The tasks colored with green in the figure are executed on specific preprocessing platforms of specific AM machines.Figure 3, Main planning tasks of process planning in Multi-part production context.1261To solve these process planning tasks in multiple parts production context, the key issues should be identified. Dutta et al. [11] regarded process planning in AM as computation problems since the planning models, algorithms, optimization and decision making models require much computation as well as the reuse of production knowledge. In this paper, feature and knowledge are used to design solutions to deal with the key issue for each of the identified planning task. And an additional task, ‘Grouping/Clustering parts’, is proposed to reduce the computation for some planning tasks with a combinatorial characteristic.

And an additional task, ‘Grouping/Clustering parts’, is proposed to reduce the computation for some planning tasks with a combinatorial characteristic. The next section will introduce the development of a feature and knowledge based systematic process planning strategy.Development of the proposed strategyManufacturability analysisManufacturability analysis is the first planning step either in multiple parts production context or single part production context. The output of this task is the feedback on the availability of processing a given part by using AM processes. The main content of the analysis contains geometric analysis and non-geometric analysis. Geometric analysis refers to analyze the size of a part and key geometric features of a part. Current AM machines have limitations on the processing size. Some parts with sizes that exceed the build volume and cannot be decomposed to build would not be processed by AM processes.

Current AM machines have limitations on the processing size. Some parts with sizes that exceed the build volume and cannot be decomposed to build would not be processed by AM processes. Some parts may be filled into a build volume, but they may collide with the boundaries of the build volume when they are rotated in the build volume. This cannot guarantee the parts to be built in good orientations with acceptable production quality. Another type of geometric analysis, analysis of surface features, is more difficult. Although AM processes can build any geometric shape theoretically, they also have limitations. The limited layer thickness may have difficulty to build some shape features with very small sizes. The slicing procedure may cause some problem to build facing features (two parallel plane faces with a very small distance).

The slicing procedure may cause some problem to build facing features (two parallel plane faces with a very small distance). The complicated distortion of AM processes cannot guarantee the shape accuracy of some features like long tiny holes or cylinders or large planes, shells, etc. Therefore, to conduct the geometric analysis, data base of processing characteristics of AM processes or benchmarking results of AM processes should be provided. Apart from this, a key or problematic feature base should be constructed to help identify the possible problematic areas on a given part model. Hence, feature recognition algorithm is also required. The non-geometric analysis is mainly focusing on the analyzing of the production requirements, e.g. part quality, build time, cost etc. For executing this analysis, benchmarking results of AM processes should be provided.

part quality, build time, cost etc. For executing this analysis, benchmarking results of AM processes should be provided. Besides of the information or knowledge data base, decision tool is also needed to act as reasoning or searching during the decision making. Many former methods adopting ‘Screen’ method which eliminates alternatives according to the checking of decision attributes, e.g. size, time, cost, tensile strength, etc, one by one. This method is efficient for the geometric analysis. However, when conducting non-geometric analysis, it would miss some potential alternatives which are very close to meet the production requirements. Therefore, to prevent the missing of some potential alternatives, an integrated decision making model (MADM) [12] is adopted for the non-geometric analysis by providing the deviation extent of1262each decision attribute.

In real application, when some designs can be modified, the deviation extent evaluation result may help to dig or attract more potential production possibilities. Therefore, a functional module for this planning task is depicted in Figure 4. The interaction and control is used to modify and control the input and output as well as set the decision attributes.Figure 4, Functional module for ‘Manufacturability Analysis’.Process and manufacturing scenario selectionWhen the manufacturability analysis is finished, an evaluation result will be obtained. If a part can be processed by AM processes, then there usually a set of finite alternative manufacturing scenarios (machine, setup, material, etc.) to produce a prototype according to production requirements. Hence, the second planning task is to evaluate those alternatives and identify the optimal one according to the production requirements and preference.

Hence, the second planning task is to evaluate those alternatives and identify the optimal one according to the production requirements and preference. The input of this task is the alternatives generated by the manufacturability analysis. The output of this task is the rank of alternative, either AM process or manufacturing scenario. To conduct the evaluation, benchmarking results of AM processes and related manufacturing scenarios should be provided as information or knowledge base. Then a decision making model is necessary to generate evaluation index for the decision support. Besides, the user preference and setting of attributes are required during the evaluation. Hence, human interaction and control are inevitable. Therefore, a full solution for this task can be depicted by a function module as described in Figure 5.Figure 5, Functional module for ‘Process & Scenario Selection’.Prediction (time, cost and quality)1263Prediction is very important in AM process planning.

For upstream departments, it is used to give quotation/pricing and support the communication with clients or even help the redesign; for downstream sectors, it is useful for the optimization and decision making during the ‘micro planning’ stage. The main contents include the estimation of build time, cost and production quality. However, this task is also very difficult to accomplish. Fast, simple analogical or empirical estimation models are usually efficient for the build time and cost estimation but they suffer from the low accuracy problem. Analytical models can give more accurate estimation results but they lose efficiency since they need detailed process planning results. In this multiple parts production context, another big challenge of build time and cost estimation is how to determine the build time and cost for individual parts which are built simultaneously. For the production quality estimation, it is more complicated. Currently, there are two types of methods.

For the production quality estimation, it is more complicated. Currently, there are two types of methods. One is to use mathematical/numerical methods to compute on the geometric models or simulate the processing procedure. This type of methods needs many mathematical/numerical or geometric assumptions due to the unknown complexity of phenomenon from physical or chemical or multiple coupled fields. Hence, the accuracy cannot be guaranteed. In addition, the efficiency of this model is relatively low due to the large computation. Another type of methods is to use experimental results for constructing empirical models. These models have a better accuracy and higher efficiency. However, they need large quantity of experiment results and production knowledge. Another difficulty of this task is that there is usually no generic prediction model for all the AM processes and manufacturing scenarios. Specific models should be constructed for specific processes and scenarios.

Specific models should be constructed for specific processes and scenarios. Hence, this task is a knowledge-intensive and computation-intensive problem.To solve this task, two types of prediction methods from literature are used. To give fast prediction of build time and cost for the quotation/pricing, analogical or empirical estimation models [13, 14] are adopted. To construct different estimation models for different AM manufacturing scenarios, different production record data bases and processing specifications of these scenarios, usually stored in benchmarking result base or machine resource base, should be provided. To give accurate estimation of built time and cost for individual parts in multiple parts production context, the analytical generic build time modeling method proposed in [15] is chosen. To use this method, processing specifications of each AM manufacturing scenario should be provided.

To use this method, processing specifications of each AM manufacturing scenario should be provided. To predict the production quality, knowledge based method is used to give fast prediction. To support the optimization and decision making in the ‘micro planning’ stage, parametric models are chosen, e.g. the surface roughness prediction model used in KARMA platform (http://www.femeval.es/proyectos/karma). Therefore, a full solution for this task can be obtained and described in Figure 6.Figure 6, Functional module for ‘Prediction’.1264Orientation optimizationOrientation optimization is one of the key planning tasks to guarantee the production quality of parts. When a group of parts to be produced in one build, the orientation problem becomes more complicated. To guarantee each part’s production quality, each part should be built in its optimal or near optimal orientation.

To guarantee each part’s production quality, each part should be built in its optimal or near optimal orientation. To diminish the total build time and cost of one build, an optimal combination of parts’ orientations should be found to diminish the total cross section area, total support volume or maximum build height, etc. However, for one given part, theoretically, there are possible infinite alternative build orientations. Therefore, for a group of parts in one build, there would be more possible infinite alternative orientation combinations. Even for a group of parts with finite alternative orientation sets, the number of alternative combinations will have a near exponential growth as the number of parts and parts’ alternative orientations increase. As a result, a combinatorial NP-complete problem forms.

As a result, a combinatorial NP-complete problem forms. Furthermore, when more objectives (attributes or criteria), such as minimizing the overall part surface roughness, overall volumetric error, total build time, total build cost, maximizing the overall part accuracy, etc., are taken into consideration during optimization or decision making, the NP-complete problem becomes multi-objective NP-complete problem with more complexity. Although the orientation optimization problem in multiple parts production context is more complicated and difficult than that in single part production context, the main sub tasks are similar. The first one is to generate alternative orientations for individual parts within a part group. However, the methods for single part orientation optimization proposed in literature cannot be directly used for multiple parts orientation problem.

However, the methods for single part orientation optimization proposed in literature cannot be directly used for multiple parts orientation problem. Because, on the one hand, the computation would be huge if too many alternative orientations are generated for each part, especially for those searching in an infinite solution space; on the other hand, not all the alternative orientations of each part can guarantee an acceptable production result for the related part. As discussed above, the general objective of the multiple parts orientation is to guarantee each part’s production quality and at the same time to diminish the total build time, cost, overall accuracy error etc.

Therefore, the first sub task for the multiple parts orientation problem is to efficiently generate a set of practical finite alternative orientations guaranteeing the production quality for each individual part within a group, but is not to rotate the parts freely respectively by doing an exhaustive searching, which would generate invalid alternative orientations. When the first sub task is finished, the second one is aimed to search out an optimal combination of parts’ build orientations to minimize the total build time, cost and other user concerned objectives (called global objectives). For solving the first sub task efficiently, an AM feature based orientation generation method [16] is adopted. When using this method, a set of finite alternative orientations can be obtained for each part. However, there is a need to refine the alternative orientations for each part since not all of the alternative orientations are acceptable ones to guarantee the part’s production quality.

However, there is a need to refine the alternative orientations for each part since not all of the alternative orientations are acceptable ones to guarantee the part’s production quality. Impractical alternative orientations may cause invalid build orientation combinations, which cannot guarantee individual part’s production quality, for the second sub task. Hence, to ensure all the alternative combinations are valid, a filtering process is added. The integrated MADM model proposed in [12] is adopted for the filtering. To identify the optimal orientation combination from the alternative ones, the second sub task applies a modified evolutionary algorithm. Therefore, a functional module and the detailed method are depicted in Figure 7 (a) and 7 (b).1265(a), Functional module of ‘Orientation’.(b).

Therefore, a functional module and the detailed method are depicted in Figure 7 (a) and 7 (b).1265(a), Functional module of ‘Orientation’.(b). Orientation optimization method for Multi-part production.Figure 7, Functional module for ‘Orientation’ and detailed method.Work space planningIn multiple parts production context, work space planning is inevitable. Maximizing the compactness of parts is usually set as an optimizing objective when nesting or packing parts into a machine build volume. Theoretically, parts can be placed or rotated freely when nesting or packing. However, apart from the compactness, the production quality of each part should be guaranteed. Therefore, the work space planning task is coupled with orientation optimization task. Actually, orientation optimization and nesting or packing can be dealt with simultaneously. However, this is too complicated due to the combinatorial characteristic which causes expensive computation.

However, this is too complicated due to the combinatorial characteristic which causes expensive computation. Hence, orientation optimization and work space planning are processed sequentially in this paper. The output of orientation optimization is the input of the work space planning. For some AM processes that need support structure, parts can only be nested in one layer, which is a two-dimensional nesting problem. However, the two-dimensional nesting problem is different to other classical nesting problems since the parts can be rotated around three dimensions though they can only be placed in one layer. For those AM processes that do not need support structure, parts can be packed upon each other. Hence, this is a three-dimensional packing problem. To solve the two-dimensional nesting or three-dimensional packing problem, nesting algorithms should be used.

Hence, this is a three-dimensional packing problem. To solve the two-dimensional nesting or three-dimensional packing problem, nesting algorithms should be used. These algorithms are serial ones, which place part one by one in sequence, or parallel ones, which nest or pack a group of parts simultaneously. When design or select nesting or packing algorithms, the efficiency should be considered since work space planning is a type of NP-complete or NP- hard problem. The output of this task is a group of positioned parts that can be sent to downstream planning tasks, support generation, slicing and scanning-path planning. Hence, a functional module, shown in Figure 8, can be designed for this task.1266Figure 8, Functional module for ‘Work space planning’.Grouping/clustering partsAs discussed above, the orientation and work space planning tasks are combinatorial problems. The computation of optimization is usually expensive due to the large alternative combinations.

The computation of optimization is usually expensive due to the large alternative combinations. To reduce the number of combinations, a modified group technology proposed in [17] is used to form part groups or clusters or sequences. Therefore, another planning task, grouping/clustering parts, is proposed for the multiple parts production context. However, this task is coupled with the orientation and work space planning tasks since the part group may be changed during orientation optimization or work space planning. To simplify the process planning problem in this multiple parts production context, this research set the grouping/clustering task before the orientation optimization task. When doing the latter two tasks, the part group or cluster can be changed according to the generated part sequence. This task is to generate a part sequence according to their ‘similarity’ which is not only limited to the geometric aspect. Part sequence is used to form part groups or part clusters.

This task is to generate a part sequence according to their ‘similarity’ which is not only limited to the geometric aspect. Part sequence is used to form part groups or part clusters. To form part sequence, production knowledge is required to identify attributes for ‘similarity’ measuring and a ‘similarity’ measuring model is needed. Therefore, a functional module can be proposed to solve this task as shown in Figure 9.Figure 9, Functional module for ‘Grouping/Clustering parts’.When all the modules for the main planning tasks in the multiple parts production context are built, then a full systematic process planning strategy forms.

The proposed strategy can be fixed in a systematic process planning framework as described in Figure 10.1267Figure 10, AM feature and knowledge based systematic process planning framework for AM in Multi-part production context.Implementation of the proposed strategyWith the constructed process planning framework for the strategy as proposed above, CAPP systems with functional modules can be implemented. In real application context, for the planning tasks in the ‘macro planning’ level and the part grouping/clustering task in the ‘micro planning’ level, the implementation of the related functional modules depends on specific production needs and available resources. Different feature base, knowledge base, production data base, prediction model base, etc. can be used. There is no common standard or solution for all of these tasks.

Different feature base, knowledge base, production data base, prediction model base, etc. can be used. There is no common standard or solution for all of these tasks. The authors had proposed some methods to solve the planning tasks in the ‘macro level’ [12] and a modified ‘Group Technology’ [17] for the grouping/clustering task. In this paper, for the limited space, only the implementation of other two planning tasks (orientation optimization and work space planning) in the ‘micro planning’ level is presented.Implementation of the orientation optimization moduleThe general method is already shown in Figure 7. There are two main steps: a. generating practical alternative orientation sets for each part to guarantee each part’s production quality; b. searching out an optimal build orientation combination to optimize pre-set global objectives.1268For the first step, an AM feature based orientation generation method [16] is adopted.

searching out an optimal build orientation combination to optimize pre-set global objectives.1268For the first step, an AM feature based orientation generation method [16] is adopted. Then, a refining/filtrating process is applied to select those alternative orientations which can guarantee the part’s production quality. When a finite practical orientation set is generated for each part in a part group, a practical orientation space forms. The forming of the practical orientation space can be depicted by Figure 11 below.Figure 11, Identifying practical orientation space from original infinite orientation space in Step one. (Note: A: infinite orientation space; B raw alternative orientation space; C: practical alternative orientation space)When the practical orientation space is identified, the next step is to search out an optimal orientation combination to optimize related global objectives. In this paper, a modified genetic algorithm is designed for the optimization.

In this paper, a modified genetic algorithm is designed for the optimization. The implementation of this module is realized on the Matlab platform (Version R2012b).Implementation of the work space planning moduleFor work space planning module, the input is a group of oriented parts, the output of orientation optimization module, which can only be rotated around the build direction and translated on the build platform. Hence, by using the parts’ projections onto the build platform as nesting stencils, this problem can be transferred into a classical 2-Dimensional nesting problem. In this paper, a parallel 2-Dimensional method is developed. It can be depicted by Figure 12.Figure 12, A proposed parallel 2-D nesting method.1269The first step is to project all the parts onto the build platform. Then, use polygons to represent these profiles. To avoid the contact or collision of parts after being nested, a minimum distance is set to expand the polygons.

Then, use polygons to represent these profiles. To avoid the contact or collision of parts after being nested, a minimum distance is set to expand the polygons. The third step is to apply a special genetic algorithm using polyploidy chromosome to represent the position of a polygon (three parameters) for searching an optimal nesting solution with minimum total overlap area. Based on the obtained minimum overlap area, nesting decision can be done. When a minimum overlap with the value of 0 can be found, it means that the part group can be placed into a specified region without collision, and vice versa. The main reason to use the parts’ projections for nesting is that some overhangs of parts may need support structures even their base areas are smaller than the projection areas. This will avoid the placement of parts under other parts’ overhangs. Hence, the up and down surfaces of parts will not be damaged by the support structures.

This will avoid the placement of parts under other parts’ overhangs. Hence, the up and down surfaces of parts will not be damaged by the support structures. However, this would cause some waste of work space when the overhangs have an angle to the build direction with less than 45 or 30 degrees where support structures may not be needed and other part could be possibly placed under the overhangs. But, the main objective of this research in the current stage is to testify the feasibility of the proposed method in the prototype level. Hence, this situation is not considered at present. Another reason to use the projection profiles as nesting stencils is to reduce computation cost since using voxel-based method to compute the interference between 3D models will cause more computation time. To ensure all the polygons are within the specified region during movement, a compensation translation is used to move those polygons that intersect with the region boundaries.

To ensure all the polygons are within the specified region during movement, a compensation translation is used to move those polygons that intersect with the region boundaries. The last step is to place the related 3D parts into the related specified 3D region by using the obtained optimal solution’s position data.In the current stage of this research, the main focus of nesting is ‘Decision problem’ (judging whether a group of parts can be nested into a specified region), which is the base of other nesting or packing problems since solutions for bin/knapsack/strip packing can easily be devised when given a (heuristic) solution method for the ‘Decision problem’ [18]. Modified algorithms can be developed to meet the real nesting needs according to this basic problem. The work space planning module is also implemented on the Matlab platform.

Modified algorithms can be developed to meet the real nesting needs according to this basic problem. The work space planning module is also implemented on the Matlab platform. To testify the feasibility of the developed algorithms, two illustrative examples are presented respectively in the following section.Test examples for two main modulesTest example for orientation optimization moduleTo demonstrate the implementation of the proposed two-step solution for the orientation optimization module, an orientation optimization for a part group composed by sixteen parts to be manufactured by a SLA machine is presented as an example. An assumption is made that no clear user preference is given for the optimization except a general objective on minimizing the build time & cost and guaranteeing the production quality at the same time.

The part group is shown in an experimental building envelope as depicted in Figure 13.1270Figure 13, A group of parts to be oriented.The first step is to generate practical alternative orientation sets for the parts. surface roughness (R-µm), support volume (V-mm3), build height (Z-mm), build time (T-min), cost (C- euro) and the projection area onto the XOY platform (build platform) (A-cm2) are identified as decision attributes and are taken into consideration simultaneously and equally to evaluate all the raw alternative orientations. After refining/filtrating, 16 sets of practical alternative orientations are generated for the part group. One set of practical orientations for a part is presented in Figure 14. Then, the next step is to search out an optimal orientation combination to optimize the global objectives.

One set of practical orientations for a part is presented in Figure 14. Then, the next step is to search out an optimal orientation combination to optimize the global objectives. In this example, to reduce the build time & cost and improve the average production quality, five objectives, Z-max (Zmax-mm), the maximum build height of the parts; Difference of build heights (std(Z)) ; Average projection area onto the vat bottom (Aa-cm2) ; Average support volume (Va-mm3) and Average surface roughness (Ra-µm), are set as global objectives with equal weights assigned.Figure 14, Practical alternative orientation set for a part.The aspired goal used for global optimization is obtained by conducting five single optimizations for the five global objectives. It is similar to TOPSIS method. The aspired goal is composed of the five obtained optimal values of the five objectives and it is given asAspired Goal = [80.4300, 16.1420, 2.4712, 376.4075, 4.1769].

The aspired goal is composed of the five obtained optimal values of the five objectives and it is given asAspired Goal = [80.4300, 16.1420, 2.4712, 376.4075, 4.1769]. (6)With the obtained aspired goal and the design genetic algorithm, the global optimizationcan be conducted. The parameters for the designed GA are set as:Chromosome length: 16, the number of the parts;1271Population size: 400; Crossover Probability: 0.7; Mutation probability: 0.1; Generation: 1000.The optimization result is presented in Figure 15 below.(a), Optimization procedure.Five-objective optimizationOptimal solution: C = [1 3 1 3 1 1 2 1 4 1 1 1 1 2 1 2]Best value: [80.4300, 16.7977, 4.8181, 393.5944, 5.5069]Aspired goal: [80.4300, 16.1420, 2.4712, 376.4075, 4.1769](b), obtained optimal solutionFigure 15, Orientation optimization result for the 16 parts.The optimization result shows that the optimal solution is very close to the unattainable aspired goal.

Four sub-objectives have attained a good approaching to their individual optimal solutions respectively except for the fourth sub-objective, minimizing the average support volume. If more preference weight can be given to the fourth sub-objective during the Many- objective optimization procedure, the evolutionary search would provide a solution with a better value for the fourth sub-objective. However, the values for other objectives may be affected. This is normal for multi-objective optimization problems where compromise among the investigated objectives should be often made. The example has testified the availability of the proposed two-step solution for the multiple parts orientation optimization problem.

The example has testified the availability of the proposed two-step solution for the multiple parts orientation optimization problem. Infinite orientation combination space can be greatly reduced by the AM feature based1272alternative orientation generation method and an optimal orientation combination is obtained by applying an improved genetic algorithm.Test example for work space planning moduleAn assumption is made that a group of parts are already oriented and exported from the orientation module. They are displayed in Figure 16 (a). As introduced above, the first step of nesting is to get the projection profiles and set the nesting region. In this example, a square region is selected as nesting region. The left corner of the nesting region is the origin point of the global coordinate system. And the compactness is set as 0.7.

In this example, a square region is selected as nesting region. The left corner of the nesting region is the origin point of the global coordinate system. And the compactness is set as 0.7. Therefore, after projection operation, the obtained polygons and nesting region are depicted in Figure 16 (b).(a) (b)Figure 16, (a), Six parts to be nested and a specified nesting region; (b), Convex approximate polygons.The following step is to apply the designed genetic algorithm to conduct the evolutionary searching. In the genetic algorithm, a running condition is set as: if the best fitness value is more than 0.95, then jump out the iteration and check the current obtained best solution. Because the polygon used for nesting is just an approximation and expended loop to represent a part’s projection boundary.

Because the polygon used for nesting is just an approximation and expended loop to represent a part’s projection boundary. Hence, when the fitness value is big enough and even it does not equal 1 (total overlap area is 0), the obtained related solution may meet the nesting requirement that is no collision between parts exists. Certainly, more rigorous running conditions can be set, e.g. the fitness value arriving at 1, which requires no overlap exists. The operating parameters of the genetic algorithm are set as:Chromosome length: 6, the number of the parts; Population size: 200; Crossover Probability: 0.9; Mutation probability: 0.2; Generation: 1500. Step length of translation: 1mm; Step length of rotation: 1 degree.After computation, an optimization result can be obtained. The current best fitness value is found as 0.9525 at the 1000th generation (Figure 17 (a)). The total computation time is 3162.1578 seconds. The figures presented below show the nesting result.

The current best fitness value is found as 0.9525 at the 1000th generation (Figure 17 (a)). The total computation time is 3162.1578 seconds. The figures presented below show the nesting result. As depicted in Figure 17 (b), the nested polygons have a small overlap. However, there is no collision between the parts (Figure 18) due to the expanded polygons and their approximation.1273(a) (b)Figure 17, a, Optimization procedure; b, Nested polygons with tiny overlap.Figure 18, Nesting result for the six oriented parts.The result shows that the proposed nesting algorithm is feasible to solve the 2D nesting problem of the work space planning module. However, the computation time is a little long. The behind reasons may include: large compactness; small translation step length and rotation degree; poor setting of parameters or operations for the genetic algorithm, etc. As stated before, this research focuses on testifying the parallel nesting method at the prototype level.

As stated before, this research focuses on testifying the parallel nesting method at the prototype level. Hence, more research should be done to improve the computation performance. However, it has the potential to compete with other nesting methods for AM proposed in literature since the parts can be rotated at any degree and small parts can be packed into the inner open holes of other bigger parts to further improve the compactness if concave polygons with multiple loops are used to represent projections with holes. This cannot be realized by current nesting methods proposed for AM in literature since many of them are ‘legal placement’ method that does not allow the occurrence of overlap during the nesting. Therefore, more advanced computation technical methods, faster computer languages, advanced graphics algorithms etc., can be applied to improve the nesting performance and computation performance of the proposed nesting strategy.

Therefore, further research should be done to improve the performance of the proposed method when applies it in real engineering context since computation time is very important.ConclusionThis paper presents a study on process planning for AM in multiple parts production context. A feature and knowledge based systematic process planning strategy is proposed. A process planning framework is constructed and some of the main modules of a developing1274process planning system are implemented on the Matlab platform. However, this is just the initial result of the current study. Due to the complexity of process planning problem, further research should be carried out.

However, this is just the initial result of the current study. Due to the complexity of process planning problem, further research should be carried out. Future work will be conducted to investigate the construction of AM feature base, AM production knowledge base, AM process benchmarking base, prediction model base, the improvement of decision models, optimization algorithms, program codes, the systematic analysis for the interrelations between different process planning tasks, etc.References1. Gogate, A. S., & Pande, S. S., (2008), Intelligent layout planning for rapid prototyping, International Journal of Production Research, 46(20), 5607-5631.2. Marsan, A., & Dutta, D. (1997). A survey of process planning techniques for layered manufacturing. In Proceedings of the 1997 ASME Design Automation Conference. Elsevier.3. Kulkarni, P., Marsan, A., & Dutta, D. (2000). A review of process planning techniques in layered manufacturing. Rapid Prototyping Journal, 6(1), 18-35.4.

In this current research, the depth-sensing indentation creep behavior of additively manufactured Ti-6Al-4V alloy and its mechanism were studied at ambient temperature for different additive manufacturing scan direction and scan size. Indentation creep tests were conducted through a dual-stage scheme (loading followed by a constant load-holding) at different peak load of 250, 350, and 450 mN with holding time of 400 s. In addition, microstructural quantitative analyses, using optical microscopy and scanning electron microscopy, were performed. Microstructural assessments and depth-sensing creep characterizations were then used to assess processing parameter/ microstructure/ creep properties relationships for this alloy.Keywords: Instrumented indentation, Additive manufacturing, Ti-6Al-4V, Creep stress exponent.1.

IntroductionTitanium (Ti) has been a metal of interest for transportation applications and biomedical industries for its excellent physical and mechanical properties, and biocompatibility [1], [2]. Ti- 6Al-4V alloy, out of different Ti alloy, contributes for more than half of all commercial Ti applications [1]. Additive manufacturing (AM), a layer-wise material manufacturing method, seems to be a suitable replacement for the conventional manufacturing processes of the Ti-6Al-4V alloy since conventional manufacturing processes are quite challenging.

AM process differing from conventional manufacturing processes in terms of cooling rate during and thermal processing required assessing the relationship between the microstructure, processing parameters, and properties of the additively manufactured Ti-6Al-4V alloys to determine its ability to meet demands of engineering design considerations.Consistency of mechanical properties and microstructures in additive manufacturing is still a critical issue and for two-phase Ti-6Al-4V alloy mechanical properties can be significantly varied by tailored microstructure [3]. The instrumented indentation testing technique (i.e. micro/ nano-indentation) is considered a robust, convenient, and non-destructive testing method to1393examine the microstructure/ mechanical property correlation in metals at ambient (room) and elevated temperatures including time-dependent plastic deformation.

Nano/ micro-mechanical properties could be calculated from the load (P) and displacement (h) data recorded throughout the process in the form of load vs displacement curve [4], [5].In this study, creep parameters including indentation creep behavior, creep rate, creep stress exponent (n) have been analyzed at different AM scan directions and scan sizes by using depth- sensing (instrumented) indentation technique, a reliable, convenient, and non-destructive testing technique to examine the microstructure/ mechanical property correlation in metals at ambient (room) and elevated temperatures [4].2. Experimental proceduresThe material studied in this research is an AM Ti-6Al-4V alloy fabricated via the laser powder bed fusion (L-PBF) process, a well-developed AM technique [6], in both horizontal and vertical scan direction. Two rectangular ‘Big sample’ and two rectangular ‘Small sample’ were prepared for both X (horizontal) and Y (vertical) scan direction.

Two rectangular ‘Big sample’ and two rectangular ‘Small sample’ were prepared for both X (horizontal) and Y (vertical) scan direction. Samples with tracks printed parallel to flow (argon) direction were named as “Core No Post X-big” (CNPX-big) and “Core No Post X-small” (CNPX-small) and tracks printed perpendicular to flow direction were named as “Core No Post Y-big” (CNPY-big) and “Core No Post Y-small” (CNPY-small). “No post” contour exposure was done on these specimens.

“No post” contour exposure was done on these specimens. Prior to the indentation, the surfaces of the specimens were scratch free mirror-like surface finish resulting from carefully ground with a series of progressively finer sand papers followed by fine polishing.In this study, at ambient temperature (298 K) using a U9820A Keysight Nano-Indenter G200, instrumented indentation creep tests with dual stage scheme (loading followed by constant load holding) were performed using a pyramidal Berkovich diamond indenter [4] of 65.3° face angle.At a constant indentation loading rate of 10 mN/s, various peak loads of 250 mN, 350 mN,and 450 mN were applied with dwell (holding) time of 400 s as the creep time.3. Results and discussions3.1. MicrostructureAM Ti-6Al-4V is generally an α-β alloy characterized with prior β grains that produces in an epitaxial way through several layers, grain boundary α, and α lath size [7].

MicrostructureAM Ti-6Al-4V is generally an α-β alloy characterized with prior β grains that produces in an epitaxial way through several layers, grain boundary α, and α lath size [7]. Fairly high cooling rates experienced by Ti-6Al-4V alloy manufactured by L-PBF processes generates an α-β lamellar structure associated with α-phase lamellae in a β-phase matrix. Fig. 1 shows the optical microscopy images of one of the specimens studied in this paper (CNPY small).1394Fig. 1: Optical microscopy of microstructure of CNPY-small.3.2. Indentation creep behaviorIndentation load/displacement at different peak loads of 250, 350, 450 mN along with the corresponding creep rate are shown in Fig. 2.

Indentation creep behaviorIndentation load/displacement at different peak loads of 250, 350, 450 mN along with the corresponding creep rate are shown in Fig. 2. It is worth mentioning that stress distribution in the depth-sensing indentation technique is much more complex than traditional creep tests, and at low displacement the maximum shear stress beneath the indenter exceeds the yield stress of the specimen (large tri-axial stresses in the range of some GPa) leading to creep in the materials even at ambient (room) temperature [8].Fig. 2: Indentation load versus depth at different peak loads and the corresponding creep rate vs. dwell time for the CNPX-big sample. Transient and steady state-creep are observed in the creep rate data.3.3.

dwell time for the CNPX-big sample. Transient and steady state-creep are observed in the creep rate data.3.3. Creep stress exponentFigures 3 a to d shows lnindversus lnindfor all conditions studied in the present paper.Indeed, the creep stress exponent (n) predicts the creep behavior illustrating creep stability and the dominant creep mechanism during depth-sensing indentation tests [9] which were deduced fromthe slope of lnindversus lnindcurves in the steady-state creep region. Based on the calculatedslopes, in all cases and conditions, the n values were recorded values greater than 3. This shows1395that the indentation creep mechanism for the printed Ti-6Al-4V materials at ambient (room) temperature is controlled through power-law (dislocation) creep [10]. Since the temperature here is well below the recrystallization temperature of Ti (i.e.

Since the temperature here is well below the recrystallization temperature of Ti (i.e. 50%Tm), the produced results are expected as no grain boundary sliding (GBS) and/or diffusion creep are involved here. It is worth mentioning that these data extracted from the steady state creep region and not necessarily the transient region.a1396bc1397dFig. 3: Creep stress exponent (n) across different samples for (a) CNPX- big, (b) CNPX- small, (c) CNPY- big, (d) CNPY- small.4. ConclusionsThe following conclusions were drawn from the of this investigation of AM Ti-6Al-4V alloyscanned with different print parameters using instrumented indentations:Creep parameters are function of indentation load and time during constant-load holdingstage.At low displacement the maximum shear stress beneath the indenter exceeds the yieldstress of the specimenSecondary stage of creep is dominated by dislocation movement (power-law creep).

alysis of Bonding Methods for FDM-Manufactured PartsaDavid Espalin, aKarina Arcaute, aEric Anchondo, aArturo Adame, aFrancisco Medina, bRob Winker, bTerry Hoppe, aRyan WickeraThe University of Texas at El Paso, W.M. Keck Center for 3D Innovation, College of Engineering, El Paso, TX 79968-0521 bStratasys, Inc., Eden Prairie, MN 55344AbstractThe fused deposition modeling (FDM) additive manufacturing (AM) technology has been valuable for producing a variety of concept models, functional prototypes, end-use parts and manufacturing tools using a range of durable thermoplastic materials. The largest individual component that can be produced in FDM depends on the dimensions of the build chamber for the specific FDM system being used, with a maximum build chamber size available of 914 x 610 x 914 mm. This limitation is not unique to FDM as all AM systems are constrained by a build chamber.

This limitation is not unique to FDM as all AM systems are constrained by a build chamber. However, by using thermoplastic materials, individual components can be bonded together using different methods to form a single piece. Bonding can be used to help reduce building time and support material use, and also allows for the fabrication and assembly of final products larger than the build chamber. This work investigated different methods for bonding FDM-manufactured parts, including the use of five different adhesives and solvents as well as two different welding techniques (hot air welding and ultrasonic welding). The available FDM materials investigated included acrylonitrile butadiene styrene (ABSi, ABS-M30, ABS-M30i), polycarbonate (PC, PC-ABS, PC-ISO), polyphenylsulfone (PPSF), and ULTEM 9085. Bonding strengths were characterized by comparing ultimate tensile strengths at break and analyzing the mode of failure.

Bonding strengths were characterized by comparing ultimate tensile strengths at break and analyzing the mode of failure. Overall, the bonding method of hot air welding produced the strongest bond for all the materials investigated except for ULTEM 9085 for which the strongest bond was achieved with the two-part epoxy adhesive Hysol E-20HP.1. IntroductionBonding is an important process used in all fields of industry where the assembly of two substrates (or adherends) is required. There are many plastic assembly methods available for the joining of materials that include the use of fasteners, adhesives, solvents, hot gas welding, ultrasonic welding, friction welding, and induction welding. Please note that this list is not complete, rather limited to notable joining methods all of which have advantages and disadvantages that enhance or limit their applicability (Rotheiser, 1999; Strong, 2006).

Often, plastic joining methods are selected based on the materials’ characteristics, fabrication method, function, and the desired aesthetics of the assembly. For example, plastics with dissimilar properties (such as melting points) can be bonded with adhesives much easier than with hot air welding (Rotheiser, 1999).The joining of parts fabricated with traditional manufacturing technologies has been extensively studied (Baldan, 2004; Priyandarshi et al. 2007; Kinloch, 1990). As the industry is interested in achieving strong bonds, the choice for the right bonding methods for a given plastic is made after extensive experimentation, especially when bonding with adhesives since themethod is very specific to each industry and it is usually difficult to generalize from one application to another (Cognard, 2006). Little work is present that considers the joining of plastics parts fabricated with additive manufacturing (AM), specifically fused deposition modeling (FDM).

Little work is present that considers the joining of plastics parts fabricated with additive manufacturing (AM), specifically fused deposition modeling (FDM). FDM is a technology that enables users to create functional parts with complex geometries by extruding a semi-molten polymer through a small-diameter nozzle. The polymer is accurately deposited in such a fashion that two dimensional profiles are stacked and fused onto each other.This study analyzed the bonding of eight commonly used FDM materials by way of solvents, adhesives, ultrasonic spot welding, and hot air welding. Table 1 lists the FDM materials used, along with some of their properties. The following describes briefly the specifics for the bonding methods used.Table 1.

Table 1 lists the FDM materials used, along with some of their properties. The following describes briefly the specifics for the bonding methods used.Table 1. Properties of FDM materials.MaterialTensile Strength (MPa)Tensile Modulus (MPa)Heat Deflection (°C) @ 264 psiGlass Transition (°C)ABS-M3036241382108ABS-M30i36241382108ABSi37191573116PC-ABS41191796125PPSF552068189230PC-ISO57998127161PC682280127161ULTEM 9085722220153186Solvent joining methods work by applying a solvent to the plastic parts being joined. At the surface, the solvent diffuses in and melts or softens the material and once both interfaces are brought into contact, material from both parts fuse and create a weld when the solvent evaporates (Rotheiser, 1999).

On the other hand, adhesive joining uses an adhesive (in one- or two-part components, usually in liquid form, but pastes and solid films are also available) that bonds to the surface of the adherends by going through some chemical or physical change (Strong, 2006). Reactive adhesives (including cyanoacrylates and two-part epoxies and polyurethanes) undergo a chemical change to create a bond. When applied to the surface of the substrate, the adhesive material diffuses into the substrate and then forms a connecting network upon reaction of the components (Strong, 2006; Cognard, 2006). One of the disadvantages of bonding with adhesives is the safety risk associated with their use since most bonding solvents and adhesives are toxic.Preferred over other methods of bonding, ultrasonic and hot air welding methods are fast, efficient, and are not associated with environmental hazards (Rotheiser, 1999; Zhang et al,.2010).

Ultrasonic welding systems use a power supply to convert a typical 50/60 Hz voltage into high frequency (20-40 kHz) electrical energy that is used to create low amplitude mechanical vibrations in the range of 20-30 µm (Rani et al., 2009). This mechanical energy is amplified by a probe or horn and then transmitted by contact to the plastic parts usually using a tip. The high frequency vibrations are transmitted through the plastic until they reach an interface, where frictional energy is developed that melts and fuses the interface of both plastic parts (Rotheiser, 1999; Zhang et al., 2010, Suresh et al., 2007; Rani et al., 2008). Hot air welding uses a welding rod that is driven through a tool while being softened by hot air. The parts to be welded and the rod are locally melted or soften, and the welding rod fills the joints that facilitate the assemblage of the plastic parts.

The parts to be welded and the rod are locally melted or soften, and the welding rod fills the joints that facilitate the assemblage of the plastic parts. The heated air is produced by a hot gas torch, and the air coming out of the torch (welding temperature) should be controlled and set specifically to the type of material. With hot air welding, it is possible to achieve weld strengths of up to 90% of the parent material strength (Troughton, 2009).This study analyzed the joining of eight commonly used FDM materials using seven different bonding methods including solvents, adhesives and welding methods. Results from the research provide information that FDM users can utilize as a reference to appropriately select an effective bonding method that properly satisfies their desired outcomes. The following describes the methods and results of this study in more detail.2.

The following describes the methods and results of this study in more detail.2. Materials and method2.1 MaterialsEight thermoplastic materials were used including ABS-M30, ABS-M30i, ABSi, PC, PC- ISO, PC-ABS, PPSF, and ULTEM 9085 (Stratasys, Eden Prairie, MN, USA). Bonding methods included the use of various adhesives and solvents including cyanoacrylate-based instant glue (HST-4 Super T, Satellite City Inc., Newbury Park, CA, USA, referred to as Superglue), solvent adhesive (Proweld Professional Plastic Welder, Ambroid, West Swanzey, NH, USA), and two- part epoxy adhesives Hysol EA 9394 (Henkel, Bay Point, CA, USA), Hysol E-20HP (Loctite, Rocky Hill, CT, USA), and BJB TC-1614 A/B (BJB Enterprises, Inc., Tustin, CA, USA).2.2 FDM part fabricationAll test specimens were fabricated using a Fortus 400mc system (Stratasys, Inc., Eden Prairie, MN, USA) equipped with T16 tips.

Building parameters accommodated a slice height of 0.254 mm (0.010 inch), a solid-normal part interior style with contour width of 0.508 mm (0.020 inch), and a part raster width of 0.508 mm (0.020 inch). Slicing was performed with Insight 6.4.1 and all test specimens were fabricated in a side-edge orientation as shown in Figure 1a. Test specimens were designed so that their dimensions after bonding were equivalent to those of an ASTM D638 Type I specimen (Figure 1b). All test specimens to be bonded were designed with a double-butt lap joint whose dimensions are shown in Figure 1c.zytest specimensupport material(b)(a)x(c)Figure 1.

All test specimens to be bonded were designed with a double-butt lap joint whose dimensions are shown in Figure 1c.zytest specimensupport material(b)(a)x(c)Figure 1. Test specimen designs (a) side-edge orientation used during part fabrication (b) ASTM D638 Type I single- component specimen (c) double-butt lap joint used for all bonding methods.2.3 Adhesive and solvent joiningJoining with the use of adhesive/solvent was carried out by applying the adhesive/solvent to all areas of the lap joint of both parts to be joined. All two-part epoxies were first mixed following the manufacturer recommendations. Both parts were joined then clipped to a flat surface to ensure that the joints were properly aligned.

All two-part epoxies were first mixed following the manufacturer recommendations. Both parts were joined then clipped to a flat surface to ensure that the joints were properly aligned. After all the excess adhesive/solvent was removed, the adhesive/solvent was allowed to cure for at least 24 hours prior to testing.2.4 Ultrasonic spot weldingA high intensity ultrasonic processor (CPX 500, Cole-Parmer, Vernon Hills, IL, USA) was adapted to create spot welds at the center of the double-butt lap joint location. The ultrasonic processor was equipped with a CV33 converter, A08766HRN horn, and a spot welding tip (Sonics & Materials, Inc., Newton, CT, USA). This combination of equipment produced vibrations with an amplitude of 0.113 mm (0.00443 in.). Prior to joining the test specimens, a penetration study was conducted to determine the amplitude percentage and pulse duration that would create enough energy to fuse the two lap joints.

Prior to joining the test specimens, a penetration study was conducted to determine the amplitude percentage and pulse duration that would create enough energy to fuse the two lap joints. These parameters were found to be specific to each thermoplastic material. Spot welds were created by using amplitude percentages that ranged from 20% to 100% for 1, 2, and 3 second pulse durations. The resulting spot welds were imaged with a stereomicroscope (MZ16, Leica Microsystems, Wetzlar, Germany) and a camera (Retiga 2000-R, QImaging, British Columbia, Canada) to determine the depth of penetration. Desired penetration depths were in the range of 2/3 to 4/5 of the specimen thickness (~1.70 - 2.70 mm). No external forces, other than the weight of the converter, horn, and tip (10.95 N), were applied during the ultrasonic spot welding process.2.5 Hot air weldingA simple fixture, shown in Figure 2, was used to align and clamp the two parts to be joined.

The material around the lap joint was melted using a Leister Hot Jet S hot air tool (Leister Technologies LLC, Itasca, IL, USA). The lap joint was then filled by using the hot air tool and a filler rod of the same material being joined.2.6 Tensile testingAll specimens were conditioned at 23 ± 2 ºC and 50 ± 5% relative humidity for at least 40 hours prior to tensile testing. Tensile tests were conducted on an Instron 5866 sytem (Instron®, Norwood, MA, USA) following guidelines provided in ASTM D638. A 10kN load cell was utilized while employing a deformation cross-head speed of 5 mm/min. Each material and bonding method was tested using at least 5 specimens to report an average modulus of elasticity, ultimate strain, and ultimate stress. Data collected during the tensile testing was used to create stress-strain curves. The modulus of elasticity, E, was defined as the slope of the initial linear portion of the stress-strain curve.

Data collected during the tensile testing was used to create stress-strain curves. The modulus of elasticity, E, was defined as the slope of the initial linear portion of the stress-strain curve. The ultimate stress, σu, was defined as the maximum stress endured by the test specimen and the corresponding strain was defined as the ultimate stress, εu.3. Results and discussion3.1 Ultrasonic Spot Welding ParametersTable 2 summarizes the specific parameters for each material (amplitude percentage and pulse duration) that successfully fused the two lap joints. Due to the design of the tip, the ultrasonic welding created a rivet that joined the two sections of material together.hot air toolfiller rodparts being joinedalignment fixtureFigure 2. Hot air welding setup.Table 2.

Hot air welding setup.Table 2. Ultrasonic welding parameters used to produce spot welds in tensile test specimens.MaterialAmplitude (%)Pulse Duration (s)Penetration Depth (mm)ABS-M306022.72ABS-M30i6022.15ABSi7022.63PC-ABS7022.50PPSF9032.52PC-ISO9022.50PC9022.43ULTEM 90858032.733.2 Tensile TestingEach of the eight FDM materials were bonded using seven different bonding methods. Sample groups of at least 5 specimens were produced. In addition, one sample group was composed of single-component (non-bonded) specimens for each material that was bonded. The following sections present the results (mean modulus of elasticity, mean ultimate stress, and mean ultimate strain) obtained from the tensile tests.3.2.1 Tensile Ultimate StressThe sample groups comprised of the single-component specimens were used as benchmarks for all bonded sample groups.

Materials were grouped and graphed according to the mean ultimate stress associated with the single-component groups. Figure 3 contains Group A which is comprised of the materials ABS-M30, ABS-M30i, ABSi, PC-ABS, and PPSF whose single-component specimens exhibited a mean ultimate stress between 30 and 45 MPa. Group B consists of the remaining materials, PC-ISO, PC, and ULTEM 9085, that exhibited a mean ultimate stress in the range of 65-75 MPa as shown in Figure 4. The data shown in the graphs represents average ± one standard deviation.Table 3 presents the performance percentage of each bonding method when used on each material.

Ultimate stress results for FDM materials in Group A. All data represents average ± one standard deviation.For all materials, the BJB TC-1614 A/B adhesive exhibited a low performance % when compared to the single-component samples of the same material. The Hysol EA 9394 two-part epoxy adhesive exhibited similar results. All sample groups bonded with the two-part adhesive Hysol E-20HP yielded a performance % range of 0-39% with the exception of ULTEM 9085 whose performance % range was 40-59%. All ABS blend materials joined with the solvent adhesive Proweld demonstrated a performance % range of 40-59%. All other materials bonded with Proweld achieved a performance % range of 20-39%.

All ABS blend materials joined with the solvent adhesive Proweld demonstrated a performance % range of 40-59%. All other materials bonded with Proweld achieved a performance % range of 20-39%. Similarly, all ABS blend materials bonded with Superglue (cyanoacrylate-based adhesive) yielded a performance % range of 20- 39% and all other materials ranged between 0-20%.All materials joined with ultrasonic spot welding exhibited a performance % of 20-39% with the exception of PC-ISO and ULTEM 9085 (performance % of 19 and 15%, respectively). The bond method that performed the best was hot air welding with a performance % of 40-100% on most materials. Materials whose glass transition temperatures were relatively lower (108- 161ºC) performed best: ABS-M30, ABS-M30i, ABSi, PC-ABS, PC-ISO, and PC.

Materials whose glass transition temperatures were relatively lower (108- 161ºC) performed best: ABS-M30, ABS-M30i, ABSi, PC-ABS, PC-ISO, and PC. In contrast, materials with elevated glass transition temperatures (180-230ºC) such as PPSF and ULTEM 9085 only yielded a performance % of 20-39% when bonded with hot air welding.8070) a P M60(s s e r t s e t a m504030PC-ISOPCULTEM 9085i t lU20100Single ComponentSuperglueProweldHysol EA 9394 Hysol E-20HPBJB TC-1614 A/BUltrasonic Spot WeldHot Air WeldFigure 4. Ultimate stress results for FDM materials in Group B. All data represents average ± one standard deviation.Table 3.

Two types of failure modes (a) Failure of the adhesive as demonstrated in a PC-ABS test specimen bonded using the two-part epoxy Hysol EA 9394. Note that the two sections forming the specimen were separated without damage to the material, and the adhesive peeled off remaining evenly distributed in the lap joint of each section.(b) Failure of the material as demonstrated in a ULTEM 9085 test specimen bonded with Hysol EA 9394. Note that the fracture occurred within the material breaking completely the lap joint of the bottom section.3.2.2 Tensile Ultimate StrainFigure 6 illustrates the results obtained for the mean ultimate strain (εu) for each material of Group A. Mean ultimate strain for each material of Group B can be found in Figure 7. In general, the hot air welded sample groups closely matched the single-component samples.

Mean ultimate strain for each material of Group B can be found in Figure 7. In general, the hot air welded sample groups closely matched the single-component samples. All other joining methods differed considerably from the single-component samples.3.2.3 Modulus of ElasticityFor all materials, the stiffness (Figure 8 and 9) associated with the hot air welding method closely approximated that of a single-component. In general, bonding with adhesive, solvent, and ultrasonic welding resulted in a decrease of stiffness. Other than with hot air welding, few notable equivalents in stiffness were observed.

In general, bonding with adhesive, solvent, and ultrasonic welding resulted in a decrease of stiffness. Other than with hot air welding, few notable equivalents in stiffness were observed. These combinations included PC-ABS, PC, and ULTEM 9085 when joined with Hysol EA 9394; PC and ULTEM 9085 when joined with Hysol E-20HP; and ULTEM 9085 when joined with BJB TC-1614 A/B.0.1)m m m m/0.090.080.07ABS-M30ABS-M30i( n i a r t s e t a m0.060.050.040.03ABSiPC-ABSi t lU0.020.01PPSF0Single ComponentSuperglueProweldHysol EA 9394 Hysol E-20HPBJB TC-1614 A/BUltrasonic Spot WeldHot Air WeldFigure 6. Ultimate strain results for FDM materials of Group A.

Modulus of elasticity for FDM materials of Group A. All data represents average ± one standard deviation.1000) a P M900800( y t i c i t s a l e700600500PC-ISOf o s u l u d o M400300200100PCULTEM 90850Single ComponentSuperglueProweldHysol EA 9394 Hysol E-20HPBJB TC-1614 A/BUltrasonic Spot WeldHot Air WeldFigure 9. Modulus of elasticity for FDM materials of Group B. All data represents average ± one standard deviation.4. ConclusionsBonding methods play a crucial role in the joining of parts whether manufactured with traditional or non-traditional methods. Parts fabricated with FDM often are limited in size since FDM systems have building chamber limitations. Consequently, over-sized parts must be fabricated in smaller components and assembled as a post process. The joining method is dependent on the characteristics of the FDM-manufactured parts.

Consequently, over-sized parts must be fabricated in smaller components and assembled as a post process. The joining method is dependent on the characteristics of the FDM-manufactured parts. Results from this study showed that certain materials have a better compatibility with specific bonding methods. For example, materials containing ABS and joined with superglue or Proweld bonded better than materials that were not an ABS blend. Additionally, hot air welding generally performed better than all the bonding methods analyzed in this study.This study provides information (Table 3) that FDM users can utilize as a reference to appropriately select an effective bonding method that properly satisfies their requirements. Adhesive and solvent bonding methods are suggested when one values aesthetic results and is not concerned with mechanical property performance. Please note that using adhesives andsolvents can become time consuming as curing times must be taken into consideration.

Please note that using adhesives andsolvents can become time consuming as curing times must be taken into consideration. If one desires to avoid curing times, then ultrasonic welding or hot air welding is recommended. While ultrasonic welding is fast, economical, and relatively easy, the results of this study showed that the assemblies created with ultrasonic welding did not exhibit high mechanical property performance. Conversely, hot air welding provided the advantages of ultrasonic welding while maintaining high mechanical property performance. The use of bonding methods in AM can help reduce building time and support material use, and also allows for the fabrication and assembling of final products larger than the build chamber.AcknowledgementsThe research presented here was performed at the University of Texas at El Paso in the W.M. Keck Center for 3D Innovation.

Keck Center for 3D Innovation. Support was provided, in part, by the University of Texas System Louis Stokes Alliance for Minority Participation Program under grant NSF-HRD- 0703584. The authors are grateful for the assistance of Eric Jerold Strickler on various aspects of the project. The authors are also grateful to Stratasys, Inc. for participating in the research and providing the test specimens used in this study.ReferencesBaldan, Adhesively-bonded Joints in Metallic Alloys, Polymers and Composite Materials: Mechanical and Environmental Durability Performance 2004Kinloch, A. J. Adhesion and Adhesives, Science and Technology. Cambridge, UK Cambridge University Press. 1990Rani, M. R., K. Prakasan, and R. Rudramoorthy. "Study of Different Joints for Ultrasonic Welding of Semicrystalline Polymers." Experimental Techniques, 2009: 36-42.Rani, M. R., K. Prakasan, and R. Rudramoorthy.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperANALYSIS OF BUILD DIRECTION IN DEPOSITION-BASED ADDITIVE MANUFACTURING OF OVERHANG STRUCTURESMohammed A. Isa*, Ismail Enes Yiğit*, Ismail Lazoglu*Koc University, Manufacturing and Automation Research Center, Mechanical Engineering Department, Rumeli Feneri Yolu, Sariyer, Istanbul 34450, TurkeyAbstracto f complex parts. However, the requirement for each layer to be structurally supported can make parts with overhangs hard to produce without alterations to the parts. This work proposes using multi-axis additive manufacturing to fabricate and analyze freeform overhangs such as bridge structures. Multi-axis AM allows reorientation of the build direction so that overhangs can be 3D printed. Consequently, decision on the build orientation is necessary and its result should be analyzed.

Multi-axis AM allows reorientation of the build direction so that overhangs can be 3D printed. Consequently, decision on the build orientation is necessary and its result should be analyzed. The effect of the AM build face directions will be studied. A Rhinoceros® plugin is direction with respect to the overhang’s local sur designed to generate the path of the multi-axis AM for the unsupported components like roofs, bridges and protrusions. The effects of the build direction on the surface quality and deformation of the components are studied.Additive manufacturing (AM) has gainedr epute asa direct method of fabricationIntroductionAdditive manufacturing (or 3D printing) became a popular rapid prototyping method since its beginningin the 1980s. It has since then expanded in application to various fields in engineering and sciences. It is now prevalent in the rapid prototyping, biomedical, construction and multi-functional applications.

It has since then expanded in application to various fields in engineering and sciences. It is now prevalent in the rapid prototyping, biomedical, construction and multi-functional applications. The most popular AM method is the fused filament fabrication (FFF) where molten material is deposited in layers to construct a part directly from a digital model. The layers are commonly extracted and filled from a tessellation model called StereoLithography (STL). This format represents an approximation of the actual analytical model using triangular surface meshes. The inaccuracy in STL files are usually negligible and do not affect the ability of AM to fabricate complex parts. Despite the complexity of parts obtainable by AM compared to other traditional manufacturing methods, the requirement for each layer to be supported limits the types of features realizable. Parts with overhanging structures like branch, roof and bridge components may not be directly manufacturable by AM.

Parts with overhanging structures like branch, roof and bridge components may not be directly manufacturable by AM. In such cases, preprocessing is commonly carried out by addition of support features to model. The additional support adversely affects production time, efficiency and quality of the part[1].An alternative way of 3D printing overhanging features is by reorientation of the part. The reorientationis aimed at repositioning the part so that a feature can be 3D printed support-less. To evaluate the amount of tilting required, it is relevant to know the overhang angle. The overhang angle is the angle by which a layer contour deviates from the vertical build direction. Features that deviate from the vertical by an angle greater than a tentative angle of 450 are usually required to be supported [2–4]. This critical overhang angle serves as the criterion to use in deciding if a part can be additively manufactured.

This critical overhang angle serves as the criterion to use in deciding if a part can be additively manufactured. By reorienting a part, a new equivalent overhang angle below the critical angle can be obtained. However, such reconfiguration needs to be carried out on a multi-axis AM(MAAM) system because many complex parts do not have any orientation where overhangs can be eliminated. MAAM is not a new concept since it has been utilized in manufacturing. In machining, it improves machining efficiency of freeform parts. Even though AM is already compliable with freeform parts, multi-axis benefits has been demonstrated in overhang manufacturing [5]. By determination of part orientation for every overhanging portion in STL, partitioning method has developed to fabricated in stages at various suitable part orientation [6].There are few studies on application of multi-axis positioning in additive manufacturing. A robot arm platform has been used for AM using multi-plane toolpaths [7].

A robot arm platform has been used for AM using multi-plane toolpaths [7]. Non-planar slicing paths has been developed for1637MAAM robot to reduce need for supports and improve production time [8]. Recently, A new slicing method termed inclined layer printing has been developed to 3D print overhang features in uniaxial AM systems[9]. The use of an inclined surface allows AM of overhangs beyond the critical slant angle. However, the pressure of compaction during deposition is less when the deposition layer is not perpendicular to the build direction. To reposition the relative build direction normal to the sliced layer, MAAM can be introduced. Research on MAAM is multifaceted comprising of design, control, process planning, path planning and analysis of manufactured parts. Most of these categories occupy an entire scope of research.

Research on MAAM is multifaceted comprising of design, control, process planning, path planning and analysis of manufactured parts. Most of these categories occupy an entire scope of research. Hence this paper will focus on the analysis of overhang features produced by AM.Multi-axis AM using 2D build platform A 5-axis AM system has been developed to facilitate variable axis of fabrication. As stated in theprevious section, MAAM can be used to change part orientation to avoid unsupported features in a model. Figure 1 shows the constructed system where a motorized build platform is attached to a 3D printer designed from Delta design concept[10]. The arms carry the print-head, which distributes material melted from the filament. The build table configuration is defined by the anglesand.𝜃𝜃𝜙𝜙Figure 1:a) The designed 5-axis additive manufacturing system.

The build table configuration is defined by the anglesand.𝜃𝜃𝜙𝜙Figure 1:a) The designed 5-axis additive manufacturing system. b) Reference frames of the 3D printer and the 2D base platform where 5-Axis relative motion between the frames is facilitated1638An initial stage in fabrication of a part is the 3D printing path planning. The path is planned within theworkpiece frame of reference. Since the part manufactured moves with the build table, it is necessary to transform each point in the part’s path to a fixed frame P. Equation 1 provides the transformation relation from the workpiece frame W to the 3D printer frame P. Where the relative position of workpiece origin respect to the printer originwithis.𝑂𝑂𝑃𝑃〈𝑤𝑤𝑥𝑥, 𝑤𝑤𝑦𝑦, 𝑤𝑤𝑧𝑧〉(1)𝑤𝑤𝑥𝑥 𝑤𝑤𝑦𝑦 𝑤𝑤𝑧𝑧𝑥𝑥 − 𝑤𝑤𝑥𝑥 𝑦𝑦 − 𝑤𝑤𝑦𝑦 𝑧𝑧 − 𝑤𝑤𝑧𝑧cos (𝜙𝜙) 0 −sin (𝜙𝜙) 1 sin (𝜙𝜙) 0cos(𝜃𝜃) − sin(𝜃𝜃) 0 sin(𝜃𝜃) 0 1 0𝑥𝑥 𝑦𝑦 � 𝑧𝑧 The transformed path in P frame is send to the printer control system where actuation takes place.

Vanek et. al[11] and Mao et.al[12] sampled points at such regions to generate support structures. In horizontal AM case where build direction remains fixed, studies have been carried out to either modify the desired part to reduce support structure or optimize the needed support structures.By changing general designed model orientation, using multi-objective optimization, Jibin [13] andPandey et. al.[14] obtained improved part orientation with respect to surface quality, support volume and production time. Optimal direction for simple models has also been studied to minimize geometric dimensioning and tolerancing errors[15]. Recent studies in design for AM have suggested inclusion of support minimization constraint at part design stage using topology optimization[16,17]. Hu et. al. investigated shape optimization and transformation of models to reduce supports. The design of support structure itself has also been studied.

Hu et. al. investigated shape optimization and transformation of models to reduce supports. The design of support structure itself has also been studied. The goal is to use the best part orientation and generate the minimum amount of support since fully self-supporting part cannot be guaranteed on a uniaxial AM system.Manufacturing and analysis of overhang structuresWith the ability to change part orientation in the making, several types of overhang structures arefabricated and analyzed in this section. The instances where the outlines of a model slants significantly away from the build direction usually results in rough surface. This is attributed to the fact that the new line of deposited material on the outline may be entirely or partially unsupported. Figure 2 shows the overhang angle , and the print platform tilt angle. Free overhangs (Figure 2b), bridges (Figure 2c) and roof, slicing anglecomponents are studied in this paper.

Figure 2 shows the overhang angle , and the print platform tilt angle. Free overhangs (Figure 2b), bridges (Figure 2c) and roof, slicing anglecomponents are studied in this paper. 𝜃𝜃0 𝜃𝜃𝑠𝑠 The lack of research and tools in non-planar slicing has been acknowledged[18,19]. Hence most MAAM systems are used to fabricated parts using planer slicers at different build orientation. The overhangs need to be identified as explained in the previous section. Moreover, for MAAM, the model is required to be partitioned after identification of support requirement. Lee et. al.[20] used adjacent layer features classification to label features and partition them into buildable and unbuildable regions. In this study, the overhang features are predesigned separately to investigate them. A common two-pass printing path is used to fabricate the overhang structures. The path obtained directly from the analytical model of an overhang part using a developed Rhino plugin.

A common two-pass printing path is used to fabricate the overhang structures. The path obtained directly from the analytical model of an overhang part using a developed Rhino plugin. A solid overhang model, in STEP or IGES file format, is first selected. After the starting layer is specified, the plugin generates the 3D printing paths.The three types of overhang parts considered can be constructed in similar matter. It is observed thatfabrication of bridge components is harder than roof components because roofs are supported at the sides. The side supports on roofs makes deposition of each raster easier due to the supports on both ends. Also, the free1639hanging component can be manufactured using similar design with the bridge, with elimination of the end wall. Hence, the analysis of overhang structures is carried out using the model in Figure 2a. Small variation in the design is used to study roof and free-hanging parts.

Hence, the analysis of overhang structures is carried out using the model in Figure 2a. Small variation in the design is used to study roof and free-hanging parts. The free-hanging part does not require collision check like the bridge component. The filament used is Colorfabb PETG filament of diameter 1.75. Fabrication of the base region of the designed model that holds the overhang structures is fabricated using normal horizontal AM using Simplify3D® for g-code generation.Figure 2:Demonstration of a) the slice path and tilt angle used in manufacturing thin overhang components, b) overhang angle definition for an overhang, c) slice angle which is equated with the tilt angleThe bridge component introduced has an overhang angle of 900. For bridge length greater than 10mm[21], there should be added supports to prevent deposition on air. Upon tilting the bridge, the effective overhang ( . The tilt angle is chosen to be equal to the slice angle.

Upon tilting the bridge, the effective overhang ( . The tilt angle is chosen to be equal to the slice angle. For slice angles ranging from 250 to 700, the bridge component is manufactured.) can be decreased𝜃𝜃𝑒𝑒𝜃𝜃𝑒𝑒 = 𝜃𝜃0 − 𝜃𝜃𝑡𝑡Figures 3a and 3b shows an overhang fabrication with high and low slice angles respectively. Whiledeposited material in Figure 3a rests on a rigid layer, the one in Figure 3b hardly touches the bridge with insufficient support. Another issue with low slice angle is the blade-like ends. The sharp end of the low slice angle bridge makes the supporting structure less rigid and less heat dissipative. Hence, for small slice angles, the tip of the bridge during fabrication is fragile and neither supports the deposited material nor dissipate the heat away from the nozzle. The top surface of the component is analyzed using contact type CMM with Renishaw MH 20 probe of 1mm diameter, as shown in Figure 3c.

The top surface of the component is analyzed using contact type CMM with Renishaw MH 20 probe of 1mm diameter, as shown in Figure 3c. It is used to study the form of the bridge surface at various slice angle.1640Figure 3:Fabrication of a) part with high slice angle, b) low slice angle using 5-Axis AM 3D printer. c) CMM measurement of surface deviation of sample partsThe CMM results of four slice angles are shown in Figure 4. Around 150 points are sampled from each surface and the deviations from the expected planar surface are analyzed. From the results found, it is observed that it is not only the surface roughness that is affected by slice angle[9], but also the overall shape of the surface. The complete results of average deviations from the plane at different slice angle is given in Figure 5.1641Figure 4: Error distribution at sample points on the top surface of the fabricated overhangs for various tilt angles.

At the sides where the nozzle turns, there are extra accumulation resulting in elevation. The range of deviation in error shrinks when the tilt angle is increasedSome similarities are observed in Figure 4. There a is common dent at the center of the bridges and sides of the bridges are raised. Since the bridge parts are thin, the motion of the nozzle at the sides carries unsolidified material to form a fold shown in Figure 6. This is what accounts for the raised sides. Figure 6 also shows that the folding phenomenon also decreases as the tilt angle is increased. It can also be observed that this behavior does not exist on the roof sample because of the side walls that increase heat dissipation at the side corners of the roof. For the roof sample that is fabricated at slice angle of 400, the average deviation is 0.051mm. This is less than all the bridge samples measured.1642Figure 5:Plot of average planar deviation on bridge components.

This is less than all the bridge samples measured.1642Figure 5:Plot of average planar deviation on bridge components. These are the errors of the manufactured surface measured with the desired CAD model of the bridging overhangsThe shape deviation plot in Figure 5 is obtained using repeated measurements from two samples of each part at the slice angles. The deviations of the fabricated surface from the ideal model are analyzed in MATLAB.The higher the slice angle (which is set to match the tilt angle), the lower the effective overhang angle,which is expected to produce better surface. The tilt angle serves the purpose of changing the part’s altering the build direction on the part. From the work of Zhao et.al.[9], where the slice angle and overhang angles are studied, there are combinations of the slice and overhang angles that lead to failure in part fabrication. In the paper, part failures are classified into two types.

In the paper, part failures are classified into two types. For slice angles that deviate too much from the overhang angle, the mode of failure is collapse failure. The other failure mode is the adhesion failure where slice planes deviate from the build direction. The adhesion failure is completely resolved by multi-axis AM in this paper. The collapse failure can be avoided by choosing appropriate slice angles. Introduction of tilt angle in this article expands the possible fabrication ranges of the slice and overhang angles.1643Figure 6: Image of sample overhang parts showing bridge, free hanging and roof components. Marked regions show the possible occurrence of folds. As the tilt angle is increased, the build building condition of the overhang improves, and better parts are producedThe path planning script enables manufacturing of freeform overhang structures as well. This is becausethe AM toolpaths are obtained from the actual overhang geometry surface.

This is becausethe AM toolpaths are obtained from the actual overhang geometry surface. For a curved surface, a new angle is introduced as indicated in Figure 7a. The angle defines the tool orientation. It is chosen and fixed from the local surface normal direction while maintaining perpendicularity with the relative tool velocity vector. An identical two-pass paths are generated in each layer using 3D curves on the overhang surfaces. Figures 7b and c show error measurement and fabrication of a freeform roof component respectively. The surface normal relative tool direction is 400. A flat roof fabricated with the same angle in Figure 7d is shown to have less error in Figure 7e.1644𝜃𝜃Figure 7: a) A freeform model where surface is used to generate AM toolpath with fixed tool orientation from the local normal b) Error of produced part surface from the designed model in a.

c)The freeform roof in a is being fabricated on the 5-Axis AM system with 400 tool direction from the normal d) A flat roofing overhang components manufactured at 400 e) Error distribution on the flat surface of the roof in d.ConclusionThe effects of changes in build direction on the shape of overhang features are studied in this paper. Theslice angle is matched with each build direction to study the effect on the shape of overhangs. For an overhang angle of 900, the inclined layer slicing can only be used for slice angles between 40 to 50 degrees. However, using a changing build direction, thin overhangs are produced with slice angles from 25 degrees and above. It is shown that with the aid of tilting, slice angles from 250 angle and above is obtainable using a PETG filament.Some factors that can affect the test like overhang thickness, printing temperature, filament feed rate and speed are kept constant for all the tests.

alysis of Droplet Train/Moving Substrate Interactionsin Ink-Jetting ProcessesS. Fathi a,*, P. M. Dickens a, R. J. M. Hague a, K. Khodabakhshi b, M. Gilbert ba Rapid Manufacturing Research Group Wolfson School of Mechanical and Manufacturing Engineering b Department of Materials Loughborough University, Loughborough, Leicestershire, LE11 3TU, UKReviewed, submitted September 10, 2008Abstract Ink-jetting technology has been applied to several processes in solid free-form fabrication (SFF) wherein droplets impinge onto a substrate to deposit the build material. Droplet impact behaviour on a surface has been the interest of many researchers; however, few studies have been undertaken to investigate the interaction of droplets with the moving substrate. This paper reports the impact behaviour of the droplets jetted at different frequencies onto a substrate moving over a range of velocities.

This paper reports the impact behaviour of the droplets jetted at different frequencies onto a substrate moving over a range of velocities. The phenomena associated with the interaction were classified into three main regimes.1 Introduction Several processes have been developed in the area of solid free-form fabrication (SFF) using the concept of material deposition [1]. Jetting is gaining more interest these days as a result of latest developments in the jetting heads. The technique is capable of depositing pico- litre sized droplets onto a substrate. This can provide of a high resolution and good final surface quality with a completely dense structure. These advantages can be applied in a layer- wise process where a functional part is made for actual use instead of a prototype.

These advantages can be applied in a layer- wise process where a functional part is made for actual use instead of a prototype. The authors are investigating such a possibility at Loughborough University by jetting two mixtures of caprolactam with initiator and catalyst to polymerise nylon on the substrate under appropriate conditions (e.g. temperature and inert gas). The concept of the process is similar to cast nylon but has lots of challenging issues with jetting, droplet placement, spreading and mixing the droplets of the two mixtures jetted separately by a drop on demand (DoD) mode jetting head. In addition, a rapid polymerisation of each layer is required which can be achieved by precise selection of mixture composition, droplet deposition and mixing method. Droplet formation characterisation and its impact behaviour on a moving substrate will be two of the main challenges in the research work.

Droplet formation characterisation and its impact behaviour on a moving substrate will be two of the main challenges in the research work. A good understanding and control of these two phenomena is required as prerequisites of the research work. Initial work studied continuous ink-jetting (CIJ) a diluted solution of a biodegradable UV (ultra-violet) sensitive resin onto a moving substrate. The achievements of this study are reported in this paper.The impact of a droplet onto a substrate accompanies a number of issues such as mechanical deformation which depends on the droplet’s physical properties and the surface conditions [2]. The droplet may stick to the surface, bounce off or splash and split into smaller droplets [3]. In the case of a stationary substrate, a droplet will spread radially to form a liquid layer [4]. In addition, depending on droplet kinetics and the impinging surface, a crown may be formed [4, 5].

In the case of a stationary substrate, a droplet will spread radially to form a liquid layer [4]. In addition, depending on droplet kinetics and the impinging surface, a crown may be formed [4, 5]. However, in the case of a moving substrate, the droplet spreading seems to be biased by the substrate motion.Corresponding Author: S.Fathi@lboro.ac.ukThere are several references reviewed in the literature describing droplet impact behaviour on a stationary surface [4-7]. A lack of literature was found in droplet impact behaviour on a moving surface, especially in SFF area. Where droplets impinge onto a stationary liquid surface, it was shown that the impact behaviour is directly dependent on the physical properties of the liquid and the frequency of the droplet train [7]. However, this paper reports the impact behaviour on a moving substrate.

However, this paper reports the impact behaviour on a moving substrate. By varying the substrate velocity and the droplet train frequency, it was planned to investigate the droplet impact behaviour quantitatively.2 Experiments A solution of isopropanol (Propanol-2 PUR from Acros Organics) and a bio- degradable resin developed for TNO Science and Industry was applied in this study. All the experiments were undertaken at a controlled room temperature of 20°C. Table 1 shows the physical properties of the solution. A contact angle measuring instrument (OCR20 from DataPhysics Instruments GmbH) was used to measure surface tension of the solution applying the pendent drop method and contact angle on a simple microscope glass slide (Lames Porte Object from Menzel Glaser GmbH) applying the sessile drop method. The dynamic viscosity of the materials was measured using a standard rotational viscometer (VT500 from Haake).

The dynamic viscosity of the materials was measured using a standard rotational viscometer (VT500 from Haake). Figure 1 shows the result.PropertyDensity Surface Tension Contact AngleValue0.98 ± 0.01 26.4 ± 0.2 22.8 ± 2Unit g/cm3 mN.m-1 DegreeTable 1. Physical properties of the jetting material70) s . P m(6050y t i s o c s i V c i m a n y D403020100020406080100120Shear Rate (s-1)FIG 1. Dynamic viscosity of the jetting solution at 20 °CFigure 2 shows a schematic of the applied setup in the current work. Two pumps were used to feed the jetting head. The prismatic pump (S/L Easy-load II from MasterFlex) supplied the jetting material through a beaker to a dosing pump (HLPC K-120 from Knauer) known as micro-pump in this paper. A damper was used to smooth the flow rate which was set at 0.55 ml/min and pressure generated by the pumps.

A damper was used to smooth the flow rate which was set at 0.55 ml/min and pressure generated by the pumps. The fluid was consequently deposited in the form of droplets through the jetting head designed and manufactured by TNO Science and Technology. The piezoelement of the jetting head was vibrated by an amplifier (LVPZT E-501-00 from Physik Instrumente Ltd. (PI)) which received sinewave signals from a function generator (TG330 from Thurlby Thandar Instruments Ltd. (TTi)).The imaging system consisted of a high speed camera (FASTCAM APX-RS from Photron Inc.) and a light source (VIT from Everest Inc.). The light source was positioned in order to capture a shadow of the droplets and impact phenomena. A lens was used to magnify the impact location of the droplets impact. Positioning of the droplets was achieved by a 3D motion system manufactured by Aerotech Inc., where a high speed slide provided a variable speed.

Positioning of the droplets was achieved by a 3D motion system manufactured by Aerotech Inc., where a high speed slide provided a variable speed. The jetting head was set at a height of 45 mm above a 1 mm thick glass slide (OCR20 from DataPhysics Instruments GmbH) fixed to the high speed slide. Imaging of the droplet impact was achieved on 35 mm of the length of the glass substrate.FIG 2. A schematic of the setupDifferent frequencies (from 5 to 20 kHz) with fixed amplitude of 0.65 v were applied to study the effect of droplet formation on impact phenomena. The appropriate frequency for the study was found to be between 9 to 18 kHz and gave a consistent droplet formation. Two frequencies of 12 and 18 kHz were typically selected as the relatively low and high frequencies in the droplet impact study. The imaging system was set to capture impact phenomena with a rate of 45,000 frames per second (fps). Substrate velocities in the range of 0.1 to 9 m/s were applied.

The imaging system was set to capture impact phenomena with a rate of 45,000 frames per second (fps). Substrate velocities in the range of 0.1 to 9 m/s were applied. The camera was triggered by remote control and impact of the droplets on the moving substrate was captured. The images were then analysed to quantify droplet formation characteristics and the impact behaviour. Droplet formation characteristics such as droplet size, velocity and generation rate, and impact behaviour such as spread layer thickness were quantified using the glass slide’s thickness as a reference.3 Results and Discussions 3.1 Droplet Formation Characterisation Figure 3 shows the effect of frequency on droplet size (diameter), velocity and generation rate. By increasing the frequency, the droplet size was decreased while in contrast, the droplet generation rate increased.

By increasing the frequency, the droplet size was decreased while in contrast, the droplet generation rate increased. There was a slight variation in droplet velocity.3.2 Droplets Impact on a Moving Substrate The droplet impact rate should be taken into account in case of a train of droplets. Table 2 shows the impact energy and its rate at different frequencies. It was interesting to find that the impact energy rate of a droplet train at 12 kHz was much higher than 18 kHz even with taking the droplet impingement frequency into account.14076)m μ () r e t e m a i D( e z i St e l p o r D12010080) s /my t i c ol e Vt e lp o r D54321600691215182169121518Frequency (kHz)Frequency (kHz)(a) (b)) c e s / s t e l p o r d ( e t a R n o i t a r e n e Gs d n a s u o h T211815129t e l p o r D636912151821Frequency (kHz)(c)FIG 3.

Variation of (a) droplet size, (b) velocity and (c) generation rateFrequency (kHz) 9 12 15 18Impact Energy (nJ) 9.7 9.4 5.7 4.7Impact Energy Rate (mJ/μs) 89 105 81 84Table 2. Impact energy rate at different frequenciesFigure 4 shows a sequence (nine consequent frames) of droplets impinging on the moving substrate. The camera captured at least three images for each droplet. The sequence in Figure 4 shows impingement behaviour of the three consequent droplets to illustrate the similarity and repeatability of the impact behaviour at the chosen conditions. As can be seen in the figure, spreading droplets made a liquid layer (film) with a specific thickness. The advancing liquid layer front was a result of both the deposition process and substrate travel. A wave formed on the liquid layer shown as W1 (Wave 1) caused by impact of droplet 1 and W2 made by droplet 2 as shown in Figure 4. It is seen that the waves shape and motion are repeatable.

A wave formed on the liquid layer shown as W1 (Wave 1) caused by impact of droplet 1 and W2 made by droplet 2 as shown in Figure 4. It is seen that the waves shape and motion are repeatable. At a high substrate velocity, some un-wetted (un-covered) sections of the substrate were caused as Figure 5 shows.Figure 6 shows formation of a wave in front of the advancing liquid layer formed with a low substrate velocity. This kind of wave had a crown-like shape which could be a result of the interface at the wet/dry surface boundary (shown in the figure) of the advancing liquid layer onto the dry surface. The existence of such interface as the source of the waves is in accordance with the theory of Yarin and Weiss [6]. However, there was a periodic behaviour in wave formation in front of the advancing liquid layer as shown in Figure 6.21FIG 4. The solution’s droplets impinging on the substrate with 8 m/s velocity (Frequency 15 kHz)FIG 5.

The solution’s droplets impinging on the substrate with 8 m/s velocity (Frequency 15 kHz)FIG 5. Uncovered sections in the droplet spreading at too fast substrate (9 m/s) (Frequency 9 kHz)FIG 6. The droplets impinging on the substrate with a high speed (1 m/s) (Frequency 12 kHz)Figure 7 shows the formation, growth and advancement of a wave and finally its separation from the advancing liquid layer. This phenomenon seems to be a result of material build-up in the advancing liquid layer front. As shown in the sequence in Figure 7, the wave separated from the advancing front of the liquid layer after reaching a certain volume. This phenomenon which will be discussed further was only seen at higher substrate velocities and jetting frequencies where the impingement frequency was high (Figure 3(c)). The separation of the built-up liquid from the advancing layer front was typically repeated three times along the monitored length of the substrate (35 mm).FIG 7.

The separation of the built-up liquid from the advancing layer front was typically repeated three times along the monitored length of the substrate (35 mm).FIG 7. Solution liquid build-up wave in front of the advancing liquid layer at a high impingementfrequency (Substrate velocity 7 m/s, Frequency 18 kHz)Figure 8 and 9 show the variation of the spread layer thickness and the height of the waves formed at the front of the advancing liquid layer with varying substrate velocity at two typical jetting frequencies. Dots in the figures indicate the spread layer thickness while the height of the waves is shown by vertical lines on top of the dots to show the rising and falling nature of the waves made above the spread layer thickness.

An image inside Figure 8 and 9 indicate how the maximum height of the waves was measured in relation to the layer thickness.The spread thickness with a substrate velocity greater than 3 m/s for the two studied jetting frequencies was constant but a thicker layer was achieved at 18 kHz frequency. As Table 2 reveals, the impact energy of the droplets and the impact energy rate at 12 kHz frequency was higher than 18 kHz which helped the deposition process. This made a thinner liquid layer of 25 μm for 12 kHz compared with the 30 μm layer thickness of 18 kHz. This was mostly affected by the higher droplet size generated at lower frequency (Figure 3(a)).The maximum height of the waves had different trends in different process conditions as Figure 8 and 9 reveal. High values of maximum wave height (seen in Figure 9) at the higher frequency (18 kHz) with higher substrate velocity (>3 m/s) corresponds with the situation shown in Figure 7.

High values of maximum wave height (seen in Figure 9) at the higher frequency (18 kHz) with higher substrate velocity (>3 m/s) corresponds with the situation shown in Figure 7. The droplets impinged onto the wet substrate and made a wave at the front of liquid layer. Growth of the wave was seen up to a certain volume until it separated from the advancing front of the liquid layer. On the other hand, lower values of maximum wave height seen in Figure 8 and 9 (lower substrate velocities), corresponds with the situation shown in Figure 6 where droplets impinged onto the wet surface and made a wave without any build-up which fell quickly before impingement of the following droplets. In addition, situations shown in Figure 4 and 5, where no wave was formed in the advancing front of the liquid layer with higher substrate velocities, were demonstrated in Figure 8.FIG 8.

In addition, situations shown in Figure 4 and 5, where no wave was formed in the advancing front of the liquid layer with higher substrate velocities, were demonstrated in Figure 8.FIG 8. Variation of the spread layer dimensions (spread layer thickness and maximum wave height) formed by the droplets jetted at 12 kHz impinging onto the substrate with varying velocityFIG 9. Variation of the spread layer dimensions (spread layer thickness and maximum wave height) formed by the droplets jetted at 18 kHz impinging onto the substrate with varying velocity3.3 Droplet/Moving Substrate Interaction Regimes In spreading a droplet, the deformation process is driven by the droplet physical properties, its impact energy and the surface physical properties [5, 6]. In the case of a moving substrate, the deposition of the droplets depends also on the substrate motion. Three main droplet/moving substrate interaction regimes were found.

In the case of a moving substrate, the deposition of the droplets depends also on the substrate motion. Three main droplet/moving substrate interaction regimes were found. Figure 10 classifies the three regimes as 1, 2, and 3.FIG 10. Spreading regimes in droplet impact on a moving substrate in different process parametersRegime 1 In Regime 1, no wave is formed at the front of the liquid layer as can be seen in Figure 4 and 10. However, a capillary wave behind the front of the liquid layer was formed as Figure 4 shows. This was generally due to a high substrate velocity typically higher than 5 m/s. Figure 8 clearly shows the zone of this regime. Generally, this regime was seen with lower frequencies and higher substrate velocity and is preferred in terms of a smooth spread layer.Regime 2 Regime 2, as shown in Figure 10, exhibits a crown-like wave at the advancing front of the liquid layer.

Generally, this wave was formed with every droplet impingement or every other impingement (Figure 6). Therefore, the periodic behaviour of formation and damping of the wave is characteristic of Regime 2. This periodic behaviour could be due to the impingement frequency and the liquid layer thickness.As shown in Figure 6, the source of the crown-like wave was the wet/dry surface boundary. This was due to the difference in surface energy between the spreading solution and the dry (glass) surface. Therefore, in an impact of a droplet on a wet surface, a kinematic discontinuity is made resulting in a crown-like wave. This is similar to the splashing droplets in impingement onto a stationary liquid surface investigated by Yarin and Weiss [6]. The crown-like wave had a height almost equal to the spread layer thickness.

The crown-like wave had a height almost equal to the spread layer thickness. Regime 2 occurred during impingement of the droplets onto the substrate with lower velocity (typically less than 3 m/s) where the substrate was made wet by previously spread droplets. Although having waves in front of the liquid layer in this regime, the waves diminished just after spreading. This regime does not affect the layer surface uniformity.Regime 3 Regime 3 was the generation of large waves (as shown in Figure 10) at the front of the advancing liquid layer. As the sequence in Figure 7 shows, the characteristic of this wave was formation, growth, travel with the advancing liquid layer onto the dry substrate and finally separation from the front after reaching a certain volume (shown in Figure 9 as maximum height). This kind of wave also had a periodic behaviour and occurred typically three times along the substrate.The height of the wave of the Regime 3 was much higher than the wave in Regime 2.

This kind of wave also had a periodic behaviour and occurred typically three times along the substrate.The height of the wave of the Regime 3 was much higher than the wave in Regime 2. Comparing Figures 8 and 9, this regime was seen with higher jetting frequency (18 kHz) and higher substrate velocities. The flow rate for the two frequencies (12 kHz and 18 kHz) was the same. Thus the main difference between Regime 3 in Figure 9 (substrate velocities higher than 3 m/s) and Regime 1 in Figure 8 (substrate velocities higher than 5 m/s) is the size of the droplets and their kinetic energy. In Regime 3 the droplets are smaller and have lower kinetic energy. In Regime 1 the larger droplets that have higher kinetic energy promote greater spreading of the material. The growth process continued to a certain volume until the inertia of the built-up wave separated it from the advancing front and it merged into the spread liquid layer.

The growth process continued to a certain volume until the inertia of the built-up wave separated it from the advancing front and it merged into the spread liquid layer. The separation volume inferred from the wave’s height increased with substrate velocity as shown in Figure 9 showing the effect of the substrate velocity in Regime 3.Although the wave eventually merged with the spread layer, it may not be flattened completely before any phase change. This could lead to a bump in the surface of thefabricated layer. Therefore the surface finish of a fabricated part could suffer from this regime.4 Conclusions The deposition phenomenon of a train of droplets in different process conditions is discussed in the current work. More specifically, the effects of the droplet formation characteristics and the substrate motion were correlated with the impact behaviour by studying the droplet spreading behaviour.

More specifically, the effects of the droplet formation characteristics and the substrate motion were correlated with the impact behaviour by studying the droplet spreading behaviour. It has been shown that three different regimes occur during the impact of droplets onto a moving substrate. The most dominant factor was the substrate velocity where waves can occur at higher values by choosing an incorrect jetting parameter. When a high substrate velocity is required, lower frequencies having higher impact energy can avoid formation of large waves which are harmful for achieving a smooth spread layer surface. By careful selection of jetting conditions and substrate velocity, it is possible to obtain a uniform layer before any phase change process. The high speed movement of the substrate helps to spread the material but this needs to be controlled carefully otherwise positional resolution and surface uniformity will suffer.References [1] Gritts, A., Valero, R., Jones, C. L.

An in-situ fatigue testing setup integrated with ultrasonic transducers and a digital microscope allows for the systematic study of fatigue crack evolution in aluminum specimens. The resulting data from experimentation, characterization, and analysis are integrated to gain unprecedented insights into the evolution of fatigue failure in wrought aluminum parts. The learning from this work will be further implemented on additively manufactured specimens.KeywordsFatigue damage, crack initiation, ultrasonic test, digital microscope, signal attenuation, discrete wavelet transform.IntroductionFatigue damage is one of the ubiquitous sources of structural degradation during both nominal and off-nominal operations in engineering components.

While model-based computational methods for assessing structural damage evolution are available, difficulty in achieving the required accuracy due to inadequacy in capturing the dynamical behavior of fatigue damage at the grain level makes them difficult to solely rely on. The random distribution of microstructural flaws such as inclusions and defects produce a wide uncertainty in the crack initiation phase [1]. Such defects become even more critical in additively manufactured (AM) components which typically consist of an array of microstructural irregularities such as voids and pores that are radically different from their cast counterparts [2].

The defects coupled with microstructural anisotropy in AM-fabricated parts render the prediction of fatigue crack evolution near to intractable through computational modeling only.Most of the existing research on fatigue analysis relies on performing the experimentation till the specimen is cracked and thereafter carrying out the post-inspection of the cracked surface. Such methods fail to accurately capture the crack initiation phenomena [3]. A recent publication suggested that the lowest performing components failed at easily observable pores, but the higher600performing components were more ambiguous with their failure mechanisms. This is due to the complex interplay between microstructural irregularities and anisotropy [4]. Studying the crack evolution in the material structure, therefore, is critical for avoiding the structural damage. Crack nucleates at the grain boundaries, and then it flows along the direction of the slip band.

Crack nucleates at the grain boundaries, and then it flows along the direction of the slip band. The maximum crack initiation length is usually less than 0.25 mm [1]. In this paper, the presence of the crack initiation is determined by finding the smallest measurable crack size on the notch surface of the specimen [5, 6].Crack initiation is highly influenced by the defects which are created during the manufacturing process. In this study, aluminum alloy specimens fabricated via wrought processing are chosen to be investigated. Fatigue loads are generated by an MTS test machine (831.10 Elastomer Test system), and the fatigue initiation of the metallic structure is monitored by ultrasonic test (UT) measurements and a digital microscope. UT is considered to be one of the best methods of nondestructive testing (NDT) that can detect the crack initiation and monitor the crack propagation. The ultrasound defines as the sound waves which have a frequency above the audible range (<20 kHz).

The ultrasound defines as the sound waves which have a frequency above the audible range (<20 kHz). In UT, the ultrasonic waves diffuse inside the specimens and detect microstructure defects such as voids and small cracks. It is well known that UT is efficient to identify the onset of the crack growth since its wavelength is very short and it can be reflected due to internal defects of the materials. To increase the credibility of UT results, the digital microscope is used in conjunction with UT to monitor the crack on the specimen surface [7].The onset of the ultrasonic signal attenuation is considered to be evidence of the crack initiation. To identify the onset of signal attenuation, the discrete wavelet transform (DWT) is applied to analyze the signals of the UT [8]. DWT can decompose the signal into the approximation coefficients and detail coefficients.

DWT can decompose the signal into the approximation coefficients and detail coefficients. The approximation coefficients represent the low-frequency signals while the detail coefficients represent the high-frequency signals.The contributions of this study are summarized as follows:1. Validation of the ultrasonic testing (UT) measurements using digital microscopic data Implementation and demonstration of the discrete wavelet transform concept on UT signals data.1. Validation of the ultrasonic testing (UT) measurements using digital microscopic data Implementation and demonstration of the discrete wavelet transform concept on UT signals data.This paper is organized in five sections. Section 2 illustrates the concept and the mathematical background of the discrete wavelet transform. Section 3 describes the experimental procedure. Section 4 discusses the results from UT, DWT, and the digital microscope. Section 5 summarizes the results and presents the future work.

As shown in Figure 1, the DWT process is divided into two sections, and it shows two-channel filter banks. The first section is called analysis filter bank and the second section is called synthesis filter bank. In this study, the analysis filter section has been studied. In the analysis filter bank of Figure 1, the original s(n) is decomposed into two channels (high pass filter and low pass filter). Since the frequency range of the decomposed digital signal is half the frequency range of the original digital (s(n)), the output of each filter is down sampled by two.y0 (k) = (cid:963) (cid:1871)(cid:4666)(cid:1866)(cid:4667)(cid:1860)(cid:4666)(cid:882)(cid:481)(cid:884)(cid:1863) (cid:3398) (cid:1866)(cid:4667)(cid:3041)(cid:3)y1 (k) = (cid:963) (cid:1871)(cid:4666)(cid:1866)(cid:4667)(cid:1860)(cid:4666)(cid:883)(cid:481)(cid:884)(cid:1863) (cid:3398) (cid:1866)(cid:4667)(cid:3041)(cid:3)The level of this step is the 1st level.

We assumed that if i=0, the filter is a low pass filter, and it produces the approximation coefficients of DWT where the low-frequency components of the signal s(n) are related to the approximation coefficients. However, if i=1, the filter is considered to be a high pass filter, and it represents the detail coefficients of DWT.Analysis filter banksSynthesis filter bankss(n)Ho2G0s(n)2Figure 1. Block diagram of the analysis and synthesis filter banks of two-channel.For multilevel filter bank analysis, eq. (2) and eq.

A notch is incorporated at the edge with dimensions of (1 mm × 3.5 mm) as shown in Figure 3. The experimental apparatus, as shown in Figure 4, is used to detect the crack evolution. Specimens are subjected to tension-tension load cycles at 60 Hz, and the load is generated by the MTS test machine. The target set point of the load is 8,000 N and the amplitude load is 4,000 N where the load fluctuates between 10,000 N and 6,000 N. In UT, ultrasonic wave is injected to the metallic specimen by a piezoelectric transducer called transmitter, and the transmitted signals are detected by another piezoelectric transducer called receiver which is located on the other side of the notch. When the crack propagates, the signals start to attenuate since part of the signals are reflected and not transmitted to the receiver as shown in Figure 6.

When the crack propagates, the signals start to attenuate since part of the signals are reflected and not transmitted to the receiver as shown in Figure 6. Furthermore, a digital microscope is used in conjunction with UT to monitor the crack evolution behavior and to provide an evidence for the attenuation of the signals. The digital microscope magnification that is used in this study is 50x.603(cid:3)(cid:4666)(cid:888)(cid:4667)(cid:3)(cid:3)(cid:4666)(cid:889)(cid:4667)(cid:3)r-_po kfr' •Figure 3. Specimen geometry used for fatigue testing.Figure 4. Experimental apparatus for crack detection.Experimental results and discussionThe digital microscope starts to observe the crack on the notch surface at almost 44,939 loading cycles while the UT does not show the attenuation in the signal at the same moment as shown in Figure 5 and Figure 6. As shown in Figure 6, the global ultrasonic data shows the signal attenuation at around 48,680 loading cycles.

As shown in Figure 6, the global ultrasonic data shows the signal attenuation at around 48,680 loading cycles. Thus, the delay of crack propagation detection using global ultrasonic data is 3,741 loading cycles. The first reason for the crack delay detection in the global ultrasonic data is that the signal attenuates only when the crack is open. At the crack initiation, the crack opens at maximum load, 10,000 N, and it closes at the minimum load, 6,000 N. The tremendous amount of the data generated during UT is another reason for the delay in detection. The total number of the signals that are measured is almost 19,404 signals with each signal having 51 measured points.604To overcome the delay detection of the crack in the global ultrasonic data, the local analysis of the ultrasonic data needs to be computed instead of the global analysis. Thus, the DWT is applied to analysis the local UT data.(a)(b)Figure 5.

Thus, the DWT is applied to analysis the local UT data.(a)(b)Figure 5. Digital microscope image of the notch surface at (a) 0 loading cycle and (b) 44939 loading cycles.The ultrasonic data, (0 cycle - 58696 cycles)Here, the start of the global ultrasonic signal attenuation..... C ·a a. re100C -~ "' "U (I) ... ::, "' re E(I)500(I)L ..... ..... 0 (I) ::, re > (I) L t-50100150 ~-~-~-~~-~-~-~--~-~-~-~023456789The measured signal pointx 10Figure 6. Measured points of the ultrasonic signals, from 0 loading cycle to 58,696 loading cycles.60510 5The block diagram process of Figure 2 is illustrated in this section as follows:1. The original signal s(n) is obtained from UT, and the length of this signal is almost 566,400 points which represent the measured signal points from the beginning of the experiment until the appearance of the crack on the notch surface as shown in Figure. 7(a).2.

7(a).2. The input of the 1st level is the original signal. In the 1st of the analysis filter banks, s(n) is 1(k)) as shown in decomposed into approximation output (y0 Figure 7(b).3. The 1st level approximation output (y0 level of the analysis filter banks. y0 and detail output (y11(k)) becomes the input signal values of the 2nd 2(k))1(k) is decomposed into approximation output (y02(k)) as shown in Figures 8(a) and 8(b).4. Step 3 can be applied to the 3rd level and the 4th level of the analysis filter banks where the input signal of each level is the approximation output of the previous level, and in each level, the input signal is decomposed into new approximation output and new detail output as shown in Figures 9(a), 9(b), 10(a), and 10(b). The ultraSOflk: data, (0 cycle - 44939 cycleJIl 100 l 50 l ~ 1 .. i -60 l100~~ I .~~ I005152253,o•200005152253(a)nw rnNW,td ~ point(b)Figure 7.

(a) 3rd level approximation output, the input signal of the 4th level and (b) 4th level of the approximation coefficients and detail coefficients of DWT.As shown in Figure 9(b), The third-level wavelet decomposition of ultrasonic signals provides a clear reduction in the approximation coefficients output which is a sign of the crack initiation since the crack appears on the surface notch at 44,939 loading cycles and the approximation output of the 3rd level decreases before 44,939 loading cycles. To increase the validity of the 3rd level results, we decompose the approximation output of the 3rd level into new approximation coefficients output and new detail coefficients output. The outputs of the 4th level of the analysis filter bank emphasize the observations at the 3rd level where the approximation coefficients reduce before 44,939 loading cycles as shown in Figure 10(b).Conclusions and future workIn this study, the importance of detecting crack initiation has been described.

The use of the ultrasonic to detect the crack initiation in aluminum specimens proved the value of UT. DWT clearly identify the crack initiation by applying the multilevel signal decomposition. The third- level wavelet decomposition is suitable to determine the crack initiation where the approximation coefficients clearly identify the sudden drop of the signal. The metallic structure can be maintained when the crack is observed during the crack initiation stage, but the digital microscope detects only the crack propagation. Therefore, more advanced measurement instrument such as Alicona confocal microscope which can monitor the crack initiation needs to6071"be used. Alicona is a 3D measurement system, and it is capable to measure the surface roughness from 3D magnified image.References[1] Gupta S. Behavioral pattern identification for structural health monitoring in complex systems. 2006. [2] Basak A, Das S.

alysis of ferroelectric ceramic fabricated by binder jetting technology Sara M. Gaytan1, Monica Cadena1, Mayela Aldaz1, Edward Herderick2, Francisco Medina1, Ryan Wicker1 1 W.M. Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, TX 2Materials Group, EWI, Columbus, OHAbstractThe M-Lab system from ExOne was used to fabricate 3D structures of BaTiO3 ceramic with applications that include dielectric capacitors, sensors, and integrated circuits. For this project, layer thicknesses of 15 and 30 μm and various percentages of binder saturation were used to fabricate components from powder. An organic binding agent was utilized during the printing process and later burned out at ~600°C prior to sintering.

An organic binding agent was utilized during the printing process and later burned out at ~600°C prior to sintering. Multiple building parameters and sintering profiles were analyzed and compared in an attempt to obtain dense parts while examining shrinkage percentage variations.IntroductionBarium titanate IV [BaTiO3] (BTO) applications range from embedded capacitance in printed circuit board to piezoelectric devices like sensors, heaters and transducers (Buscaglia, et al. 2004) (Maison, et al. 2001). Other applications include semiconductors with positive temperature coefficient of resistivity (PTCR) (Chatterjee, Stojanovic and Maiti 2003), dynamic access memories (DRAM), and IR sensors (Sahoo, et al. 2007). Synthesis techniques for BTO preparation include the organo-metallic precursor method, coprecipitation/sol-gel calcination and hydrothermal methods, among others (Sahoo, et al. 2007).

2007). Synthesis techniques for BTO preparation include the organo-metallic precursor method, coprecipitation/sol-gel calcination and hydrothermal methods, among others (Sahoo, et al. 2007). Some deposition methods are vacuum evaporation, sputtering, laser ablation, hydrothermal synthesis, radio frequency plasma sputtering and sol–gel techniques which have been used for obtaining BTO thin films (Harizanov, Harizanova and Ivanova 2004). BTO ceramics can also be processed by techniques such as Spark Plasma Sintering (SPS) (Buscaglia, et al. 2004), a change that occurs when cooling through the Curie point, resulting in the involvement of stress-related transformation.Depending on its crystal structure, BTO can be paraelectric (cubic; a=4.031Å) or ferroelectric (tetragonal; a=3.994Å and c=4.038Å) (Chen, et al. 2005) (Maison, et al. 2001).

2005) (Maison, et al. 2001). While low calcination temperatures of ~700°C result in a metastable cubic phase, this can be transformed to tetragonal structures by using higher temperatures, as found by Maison, et al. where phase transformation and particle size were analyzed by varying calcination temperature (Maison, et al. 2001). It has been demonstrated that chemical homogeneity, material purity, and microstructure have an effect on the positive temperature coefficient of resistivity in BTO. For this reason, methods that provide less contamination, such as coprecipitation, polymerization and sol-gel, are being developed (Hur, et al. 1998). For this project, BTO powder was used to fabricate simple geometric figures with a binder jetting technology, the M-Lab system (ExOne, OH). These components were fabricated with different binder saturation percentages and sintered at different temperatures to obtain maximum density. XRD data show no signs of contamination after sintering.

XRD data show no signs of contamination after sintering. Part shrinkage was calculated and used to obtain accurate dimensions for dielectric constant measurements.Material, processing, and investigation methodsA total of 500 grams of BTO powder was purchased from Acros Organics through FisherScientific. As-received powder can be appreciated in Figure 1. It can be observedthat powder particles are approximately 1μm in diameter, while agglomerations can measure up to 10μm.Simple geometries of were BTO the binder fabricated by jetting process using the ExOne M-Lab system. The basic this process was to obtain three- dimensional parts bound together in a layer-by-layer fashion. ExOne technologies consist of a two-bed system, shown in the schematic in Figure 2; one is for the base powder material and the other for part fabrication. Powder layers are spreadpowderideabehindFigure 1. As received BTO powderFigure 2.

Powder layers are spreadpowderideabehindFigure 1. As received BTO powderFigure 2. Schematic of powder bed systemwith the aid of a roller while the bed is mechanically moving in the x- direction. Once the powder layer is uniformly spread, binder droplets are deposited in the specified places for part fabrication. For this project, a 2:1 ratio of powder and building deposition was used. Fabricated parts, while still in the powder bed, were placed in the oven at 195°C for 2 hours to cure the binder, then cleaned and sintered or infiltrated with a different material. These parts were previously designed by CAD and uploaded into the system as an .stl file.For this project, different building parameters were used and post-sintering density was compared. The binder saturations (which refers to the amount of binder used while building) used were 60, 75, 100 and 120%. Since layer thickness is based on powder size, a 15μm thickness was used at the start of the project.

Since layer thickness is based on powder size, a 15μm thickness was used at the start of the project. However, since samples were braking in small pieces after binder curing, the layer thickness was doubled to 30μm, regardless of powder size. Prior to building with a new powder, or even a new layer thickness of the same powder, the M- Lab system requires to calculate and input the powder packing rate before using a new layer thickness.Two hours at 195°C were used to cure the binder and the samples were cleaned before sintering. A high temperature cyclical furnace (CM, NJ, USA) was used for all sintering temperatures. To simulate non-cyclical furnace cooling profiles, conditions were decided as follows (based on a sintering temperature of 0.8Tm). Samples were heated at a rate of 10°C/min up to 600°C and held for 0.3 hours, allowing binder burnout. They were then heated up to either 1260°, 1330° or 1400°C, held for four hours, and cooled down to room temperature.

They were then heated up to either 1260°, 1330° or 1400°C, held for four hours, and cooled down to room temperature. Built samples were left inside the furnace until reaching room temperature to avoid faster cooling. Precise temperature profiles are better-represented in Figure 3.Figure 3. Temperature profiles used to simulate a regular furnaceDensity values were obtained by following ASTM B962-08 standard test method (Standards 2009) which uses the Archimides’ principle. A Mettler Toledo MS303S precision scale was used with the aid of a copper mesh to hold the samples in place. A drop of soap was dissolved in the water to reduce surface tension, and olive oil for the immersed samples (used to seal surface porosity and avoid water adsorption), and the water temperature was ~22°C. For this reason we have two types of density, which are referred to as sintered (as removed from the furnace) and impregnated (which were dipped in oil at a pressure of 7kPa for 30 min.).

For this reason we have two types of density, which are referred to as sintered (as removed from the furnace) and impregnated (which were dipped in oil at a pressure of 7kPa for 30 min.). SEM analysis (TM-1000, Hitachi, Japan) was performed using an accelerating voltage of 15kV and no sample preparation was required. XRD analysis was obtained with Cu kα radiation (D8, Bruker, Germany). Dielectric constant was obtained with an Agiles Technologies PNA-X Analyzer with a frequency capability of 10MHz to 50GHz.Results With BTO being one of the most versatile ceramics available, the ultimate goal for thisproject was to obtain dense parts that can provide a dielectric constant comparable to otherFigure 3. Density vs Sintering Temperature for different binder saturationsFigure 4. Set of magnified images a) to c) of sample fabricated with 60% binder saturation, layer thickness of 30µm and sintered at 1200°C for four hoursFigure 5.

Set of magnified images a) to c) of sample fabricated with 60% binder saturation, layer thickness of 30µm and sintered at 1200°C for four hoursFigure 5. Set of magnified images a) to c) of sample fabricated with 60% binder saturation, layer thickness of 30µm and sintered at 1260°C for four hours.binder fabrication methods. Different saturations and temperatures were used to fabricate components and density was obtained for those parts. Figure 3 shows the results obtained; each data point depicts the average density of five measured densities. It can be observed that the maximum obtained density value is 1.6 g/cm3, which is about 27% dense for oil impregnated density measurements or 1.3 g/cm3 (22% dense) for the sintered part.Both densities were calculated since it is assumed that the porosity obtained in the parts is interconnected. In addition, unsealed surface porosity absorbs water causing the density values to be higher than the true value.

In addition, unsealed surface porosity absorbs water causing the density values to be higher than the true value. For this reason, it can be observed that two types of density were calculated, sintered and impregnated as observed in Figure 3.The implied reason for the low density obtained is the lack of pressuring force while sintering. Nonetheless, binder saturation plays an important role in this matter as well. To samples better understand fabricated with 60% binder saturation, which provided the highest density obtained, and a layer thickness of 30µm were sintered at 1200°C for a total of four hours and the surface was analyzed. Figures 4a), b) and c) show a magnified sequence of images where binder has already burned out (process that occurs at ~600°C) and empty spaces were left behind. Binder droplets are 60µm in length and 20µm width.

Binder droplets are 60µm in length and 20µm width. This means that BTO grain growth must overcome the empty spaces left behind after binder burnout, and close any gaps around the grains in order to obtain denser parts.the process,Since the application of force while sintering is not a path to follow for 3D fabricated parts (since complex parts would lose temperature was the increased to obtain denser parts. According totheir shape),Figure 6. Set of magnified images a) to c) of sample fabricated with 60% binder saturation, layer thickness of 30µm and sintered at 1330°C for four hours.Figure 7. Set of magnified images a) to c) of sample fabricated with 60% binder saturation, layer thickness of 30µm and sintered at 1400°C for four hours.Gao et al., relative density and grain size increased as sintering temperature increased from 1200°C to 1300°C for 2 hours obtaining a density relatively close to 100%.

Nonetheless, their powder particle size was 35nm and after powder pressing, sintered samples were only 1 mm in thickness with a maximum grain growth of 2.2 µm (Gao, et al. 2004). In a different study, sintering temperature for powders 30nm in size, consisted of 800°C for ten hours, obtaining a density of more than 90% the theoretical density (Buscaglia, et al. 2004). According to Srinivas, the sintering behavior can be improved by having small, non-agglomerated and equiaxed particles (Subramaniam 2006). For this project, final dimensions consisted of 10.2 x 10.2 x 2.5 mm while starting powder was equiaxed and minimum agglomeration was observed. Nevertheless, density obtained was approximately one fourth of ideal.Figures 5, 6 and 7, are sequences of images that have been taken at low, medium and high magnifications for samples fabricated with 60% binder saturation and layers of 30μm in thickness.

The difference between each set of images is that samples were sintered at 1260°C, 1330°C and 1400°C, respectively. Given the fact that calculated density for both sintered and impregnated parts provided the highest values obtained for 60% binder saturation, it can be stated that sintered grains attained grain growth. This can be observed in figure 5c) where grain size values range from 25 to 40μm, and space between grains is fully enclosed. However, Figures 5a) and b) show how most of the powder did not reach any sintering stages and its size remained unchanged. Figures 6 and 7 show less unsintered powder when compared to samples sintered at 1260°C. Images 6 and 7c) show open spaces in between grains, which is not observed in Figure 5c). Densification during sintering can occur mainly in two types of mechanisms, lattice diffusion and grain boundary diffusion, where grain boundaries are the material source for both mechanisms.

Densification during sintering can occur mainly in two types of mechanisms, lattice diffusion and grain boundary diffusion, where grain boundaries are the material source for both mechanisms. However, samples sintered at 1330°C and 1400°C show empty spaces in- between the grains when compared to samples sintered at 1260°C. Solid state sintering usually occurs in three stages which are i) initial: involving necking and grain growth, ii) intermediate: resulting in shrinkage and densification, and iii) final: elimination of close pores and formation of dense microstructure (Subramaniam 2006). In our parts, we are probably only reaching the initial stage where grain growth is occurring but it is not uniform.

In our parts, we are probably only reaching the initial stage where grain growth is occurring but it is not uniform. For this reason, further research developing triggering mechanisms for proper sintering behavior of BTO fabricated by the M-Lab system needs to be executed.Since samples sintered at 1260°C provided the highest density on both sintered and impregnated measurements, our focus converged to those samples. Transition temperatures for BTO from tetragonal crystal structure to cubic is 130°C while transformation from cubic to hexagonal is 1460°C. The issue with obtaining the hexagonal crystal structure is that BTO properties stop being ferroelectric, while tetragonal structures are both polar and ferroelectric (Subramaniam 2006). XRD analysis results showed how the BTO powder obtained is mainly tetragonal with small amounts of cubic phase.

XRD analysis results showed how the BTO powder obtained is mainly tetragonal with small amounts of cubic phase. After a sintering cycle at 1260°C for four hours, it can be observed how a cubic peak in the 200 direction at 2θ=4_° and an hexagonal peak in the 110 direction are observed side-by-side (see Figure 8). According to Chen et al., solid state sintering leads to the cubic phase while cooling, and each grain formed is surrounded by neighboring grains of random orientation closely adhering at the grain boundaries (Chen, et al. 2005), which can explain the tight space between grains in Figure 5. Another study conducted by Gao et al., showed similar results regarding the tetragonal structure’s inclination to transform to a cubic phase when sintering BTO powder to 1250°C (Gao, et al. 2004).To obtain dielectric constant values with a frequency of 12.4 to 18 GHz, samples of 15.8mm in length, 7.9mm in height and 5.08mm in width had to be obtained to use Ku band.

2004).To obtain dielectric constant values with a frequency of 12.4 to 18 GHz, samples of 15.8mm in length, 7.9mm in height and 5.08mm in width had to be obtained to use Ku band. For this reason, percent shrinkage was tabulated (Figure 9), and samples were fabricated the percent shrinkage obtained after sintering. In addition, CAD files were scaled with ±1% error to accommodate for any discrepancies during sintering. It was found that samples built taking into consideration the shrinkage percentage calculated did not need any post- processing. On the other hand, samples built to accommodate for the required dimensions that were not scaled before fabrication, required some grinding prior to analysis.consideringFigure 8. XRD analysis of as-received powder and sample sintered at 1260°C for four hoursDielectric constant values ranged from 8.62 to 6.23 for frequencies ranging from 12.4 to 18GHz at room Results are shown in temperature. Figure 10.Figure 9.

Figure 10.Figure 9. Graph depicting shrinkage percentage in each direction vs sintering temperatureFigure 10. Dielectric constant vs frequencyApplications for multilayer ceramic capacitors imply the importance of accommodating thin dielectric layers and avoid the presence of big particles and pores when designing electronic devices (Chen, et al. 2003)ConclusionsBTO promises to be a material that can be fabricated with additive manufacturing. Nonetheless, further research is necessary to obtain optimum properties.

2003)ConclusionsBTO promises to be a material that can be fabricated with additive manufacturing. Nonetheless, further research is necessary to obtain optimum properties. From the preliminary results shown in this project the following conclusions can be stated:Density after sintering must be increased and future work will include increasing sintering temperature while avoiding grain size growth.XRD analysis indicates BTO fabricated by binder jetting technology holds a tetragonal structure even after sintering, providing ferroelectric properties to the ceramic parts obtained.Dielectric constant values range from 8.62 to 6.23 as the frequency increases from 12.4 to 18 GHz. Future work consists in further investigation on this value being increased by denser parts, smaller grains and the addition of different coatings.AcknowledgementsThe research presented here was performed at The University of Texas at El Paso (UTEP) within the W. M.

alysis of Fiber Attrition and Mechanical Performance in Large-Format Additive Manufacturing of Long-Fiber Reinforced Polymer CompositesAndrew Rhodes1, Tyler Smith1,2, Christian Sharpe1, Vlastimil Kunc2, Chad Duty1,21 Mechanical, Aerospace, and Biomedical Engineering, University of Tennessee, Knoxville, TN 2 Manufacturing Science Division, Oak Ridge National Laboratory, Knoxville, TNAbstractUnderstanding the residual fiber length of discontinuous fiber-reinforced thermoplastics in large-format additive manufacturing (LFAM) is of significant interest since the residual fiber length influences the mechanical properties of the final part. Currently, the attrition of long fiber reinforcement during screw-extrusion LFAM is an understudied subject, although the residual fiber length is among the most important microstructural properties of fiber-reinforced composites.

A contributing factor the lack of focus in this area is the questionability and variety of convoluted fiber measurement methods. This study evaluates the relationship between printing speed, the final fiber length, and resultant mechanical properties for 4 mm pultruded thermoplastic polyurethane (TPU) pelletized feedstock and seeks to fully document a fiber length distribution (FLD) measurement method. Samples were printed at several processing speeds and subjected to fiber length analysis and tensile testing. Fiber length analysis was performed using a combination of novel and proven techniques for improved repeatability and reliability of results. This study seeks to improve understanding in the choice of machine design, material selection, and processing parameters for optimal mechanical properties.IntroductionBackgroundThe use of reinforcing fiber is essential to large-format polymer additive manufacturing [1].

However, the length of the reinforcing fibers is significantly altered through the extrusion process, and the extent of this alteration pertaining to the processing parameters is not well understood, and this process-induced degradation of the fibrous microstructure can induce a noticeable reduction in mechanical properties.LFAM systems use a single screw plasticating extruder to rapidly deposit large quantities of pelletized material in the printing process. One of the major challenges in printing material this way is the process-induced fiber length reduction. Screw-induced plastication is a high shear process, subjecting the fibers to high stresses that bend and eventually break the fibers [2–8]. Fortunately, this process is near identical to injection molding, where there have been numerous investigations concerning the influence of processing conditions of fiber reinforced plastics, and the resulting mechanical properties of the material. Wolf et al.

Wolf et al. demonstrated that conditions which increase shear decrease average fiber length [3]. Wolf took fiber length distributions from different regions within the extruder under varying screw speed, temperature, and die diameter. Wolf found that by varying processing conditions to reduce shear (slow screw speed, higher extrusion temperature, larger die diameter) fiber length attrition is reduced. Turkovitch et al. studied the progression of fiber degradation through the length of the extruder and at different screw speeds.256Turkovitch found that fiber degradation advances during extrusion and with screw speed [9]. Yilmazer et al. examined the effects of processing conditions on fiber length and strength in injection molded samples. They reported that fiber length and tensile properties significantly decrease as shear rate is increased by modifying the screw speed and feed rate [8]. Hausnerova et al.

They reported that fiber length and tensile properties significantly decrease as shear rate is increased by modifying the screw speed and feed rate [8]. Hausnerova et al. studied fiber degradation during multiple extrusion cycles under a range of shear rates through a capillary rheometer. Hausnerova reported that high shear rates degrades fiber length, with significant damage occurring during the first extrusion cycle [10]. In summary, processing conditions have been found to affect mechanical properties through reduction or retention of the initial fiber length distribution, and all these studies rely on accurate fiber length distribution measurement.Quantifying the residual fiber length distribution is of critical importance in studying the effects of the processing parameters, which directly influence shear and thus the residual fiber length.

Additionally, accurate and reliable fiber length distribution (FLD) data are necessary to develop automated measurement tools and predictive models for fiber attrition and mechanical properties. Despite its importance, there is a lack of a defined standard, resulting in a wide variety of incomparable fiber measurement and acquisition methods that exist within academia [11]This work examines the influence of screw speed on the fiber length of a long fiber material, and tensile properties using a large-format polymer additive manufacturing system. This subject has been extensively studied in an injection molding setting, but it has not been well tested in a screw-extrusion additive manufacturing system like the Cincinnati Big Area Additive Manufacturing (BAAM) used in this study.

This study was designed considering that the BAAM extruder is akin to an injection molding extruder, and therefore the effect of processing conditions on fiber length and mechanical properties should follow those of the studies done in injection molding. Additionally, this work examines novel techniques underutilized in FLD measurement and presents improvements to these methods.MethodsPrinted SamplesThe material used in this study was pultruded 40% wt. glass fiber thermoplastic polyurethane (TPU). The pellets were of a 4 mm nominal length, but some were found to be a few dozen microns longer (~4.2 mm). The pellets were dried at 80oC for at least 4 hours to remove moisture before printing. The processing parameter varied during this study was screw speed. All other conditions were held constant. The screw speeds tested were 40, 75, and 150 RPM, as this is the range at which the BAAM can successfully print this material.

All other conditions were held constant. The screw speeds tested were 40, 75, and 150 RPM, as this is the range at which the BAAM can successfully print this material. Samples were printed at 230 oC with an 8.89 mm (0.35 inch) diameter nozzle, 7.62 mm (0.3 inch) layer height, and a 17.78 mm (0.7 inch) bead width. Each sample set was printed as a 305 mm (12 inch) x 305 mm (12 inch) x 15.24 mm (0.6 inch) plaque, and then cut to 11.5 inches to closely fit the 11.5-inch square compression mold. The plaques were compression molded at 145 oC under 5 tons for 5 minutes each. The average final thickness for the compressed plaques was 6.5 mmTensile specimens (ASTM D638 Type 1) were waterjet cut from the plaques such that the loading direction was parallel to the printed layers or deposition direction [12]. The samples were dried at 50 oC for least 48 hours and left in a desiccant chamber at 23oC for 12 hours (ASTM D618- B) prior to tensile testing.

The samples were dried at 50 oC for least 48 hours and left in a desiccant chamber at 23oC for 12 hours (ASTM D618- B) prior to tensile testing. Specimens that incurred damage prior to testing were discarded.257Tensile testing of all specimens was performed on an MTS Criterion Series, Model 45 with a 10kN load cell. The testing rate of 5 mm/min was used for a nominal strain rate of 0.5 mm/(mm min). Extension was measured using an MTS LX 500 Laser Extensometer. Specimens that fractured outside of the gage length were omitted from the data analysis. To ensure statistical significance a minimum of 5 specimens were measured for each condition.Fiber Length Analysis Method and TechniquesThe objective of the FLD measurement process is to obtain a distribution that characterizes the microstructure of the composite and can be divided into seven steps. FLD measurement is performed by firstly isolating a small, manageable portion of the composite called a coupon.

FLD measurement is performed by firstly isolating a small, manageable portion of the composite called a coupon. Second, the matrix material is removed, often by pyrolysis. Third, a subset of the fibers is selected for measurement through a process called down-sampling. Fourth, the fibers are dispersed and prepared for image capture. Fifth, a digital image is created via microscopy or a document scanner. Sixth, the image is measured either by hand or machine using an image processing software. Seventh, the data is analyzed, and potential biases induced by the prior steps are considered and amended. In the following pages, the specific FLD methods and techniques used in this study will be explained and discussed.1. Composite coupon extractionTo accurately characterize the fiber length distribution within the polymer matrix, coupons of the printed composites were cut material via a 1.25-inch interior diameter hole saw with the guide bit removed [Figure 1].

According to Kunc et al., the coupon size should at least be twice as large as the longest possible fiber, so that fibers cut during the couponing step can be excluded from measurement simply by extracting fiber from the center of the sample [13]. Future experiments complementary to this study were planned to include a maximum of 12 mm fiber, and so a hole saw with a diameter greater than twice that length (31.75 mm or 1.25 inch) was selected to sample the coupons. A circular coupon is preferrable to a rectangular one for the speed and consistency at which samples can be taken. A consistently sized coupon is required so that it closely fits its constraining container during the matrix removal step.Figure 1: Composite coupon and the plank from which it was cut2582. Constrained removal of matrix materialThe composite coupon is placed in a 1.25-inch interior diameter aluminum pipe as the constraining vessel for pyrolization or burn-off of the TPU matrix material [Figure 2].

The bottom of the pipe was covered with aluminum foil and attached loosely via a steel hose clamp to keep the foil bottom secure. The top was left open during burn- off. Constraint of the sample is necessary to roughly preserve the structure of the sample during the expansion of the matrix material and release of elastic energy stored in they were extruded. Unconstrained burn-off can result in a pile or puff of fiber, rather than a stack. The top of the sample is left unconstrained to allow for gas to escape, as well as to allow the sample to expand vertically. Burn-off was performed in-air for 4 hours at 500°C with a heat ramp of 3°C/min using a small crucible oven. Some sources reported successful pyrolization for 2 hours at 400°C, but that left significant char on the fibers.the fiber during whenFigure 2: Constraining container and the composite coupon fiber stack3.

Fiber sample isolation / down samplingThe fiber sample isolation step involves isolating a portion of fibers from thethe coupon which represent composite while excluding the damaged fibers near the coupon’s edge. This was performed with the epoxy plug method developed by Kunc et al.[13]. This method involved inserting a 22- gage hypodermic needle attached to a 1 milliliter syringe loaded with an epoxy resin through the center of the coupon and retracting the needle while dispensing epoxy at a continuous rate, resulting in a cylindrical column of epoxy resin that encapsulates fibers within the center of the coupon fiber stack. A Bondic UV curing resin was used since glass fiber can refract the UV light, resulting in significantly faster cures than with an epoxy resin.

A Bondic UV curing resin was used since glass fiber can refract the UV light, resulting in significantly faster cures than with an epoxy resin. A 3D printed disk with a center hole for the needle was used as the lid to ensure central alignment of the needle, as well as to lightly compress the fiber to prevent them from sticking to and rising as the needle retracts. The needle was inserted through the entirety of the coupon, such that the needle tip is protruding from the bottom of the aluminum foil bottom. The resin injection was conducted using a custom designed & built actuator for precise control of the needle withdrawal rate and the Figure 3: Epoxy depositorthe FLD of259volume of resin deposited [Figure 3]. The depositor is composed of common 3D printer components (Nema 17 motors, TMC2209 stepper drivers, LM8UU bearings, etc.), and an Arduino. Two independently addressable stages allow for precise customization of the resin column size.

Two independently addressable stages allow for precise customization of the resin column size. After the epoxy was deposited, a 5-watt UV flashlight was aimed at the top of the sample for approximately 30 seconds, and at the bottom hole for 30 more seconds, after which the resin would be cured enough to remove the sample from the container. Removal of the sample was performed by sliding the hose clamp off and loosening the sides of the aluminum foil bottom from the container. Then, the 3D printed disk lid was carefully held down while lifting the container up, resulting in the removal of the fiber stack from the container. The UV light was shone on the fiber stack from the cardinal angles for 30 seconds each to aid in a more consistent cure as the samples in this study were highly fiber dense (40 weight percent). In the future, use of a stronger UV lamp would likely improve penetration of the light through the fiber stack to decrease cure time.

In the future, use of a stronger UV lamp would likely improve penetration of the light through the fiber stack to decrease cure time. Additionally, it is useful to add a small amount of epoxy to the top of the plug to act as a handle. The epoxy-bound fibers are extracted by grabbing an edge of the plug with tweezers and using a small, soft paint brush to shed uncaptured fiber from the plug.Once the plugged fibers are isolated, burn-off of the resin is performed using the same program as the first burn-off, but with the maximum temperature time set to two hours. During this burn-off, the fibers are partially disentangled via the expansion of the resin. The heat ramp of 3°C/min was sufficient for this study, but slower heat ramps can improve disentanglement [11].4. Filament dispersionA well dispersed image is important in several senses. Firstly, images can potentially contain tens of thousands of fibers, and it is unreasonable and impractical for a human to measure each one.

Firstly, images can potentially contain tens of thousands of fibers, and it is unreasonable and impractical for a human to measure each one. If a fiber field is evenly dispersed, a human operator can take a subsection or subsections of the image that contains the number of fibers needed, while being fairly confident that the fibers measured are statistically representative of the entire fiber distribution. Secondly, the elimination of fiber bundles or areas of highly dense fiber and overlaps is important for measuring the distribution by hand, through one of the few free but difficult to use automated methods, or inaccessible proprietary options for automatic fiber length measurement [11,14–16].Initial experiments with fiber dispersion methods showed inconsistency and potential breakage of fiber. These included mechanical means like brushing [13,15], submersion and manual mixing in liquid [14], and ultrasound [17].It was found that a novel method introduced by Goris et al.

These included mechanical means like brushing [13,15], submersion and manual mixing in liquid [14], and ultrasound [17].It was found that a novel method introduced by Goris et al. provided a low-input, non-mechanical solution to disentangle loose fiber bundles and evenly disperse fiber upon the glass sheet of an image scanner [11]. Under this method, the fibers are dispersed within an enclosed funnel/chamber system using gentle bursts of compressed air. The turbulence causes fibers to disentangle and disperse evenly. A can of compressed air was used with light pressure on the trigger and was found to sufficiently disperse the fiber. Since there is potential for the air pressure to be too great and blow fiber back towards the user, some care is required to lightly release air from the can, and it is recommended to add several inches of space between the mouth of the funnel and the nozzle of the air canister, or apply a seal to the mouth of the funnel and air canister nozzle.

The pressurized air from the small260canister nozzle will push a larger wall of air, which is strong enough to disperse the fibers, but gentle and enough to prevent fibers from blowing back. A detailed schematic of the dispersion funnel is shown in Figure 4. The funnel was reverse engineered using Goris’s drawings of their funnel, and 3D printed. The funnel was sketched over the original and designed to fit the image scanner used in this study. It should be noted that Goris used polished stainless steel and a dry graphite lubricant to prevent fibers from sticking to the surface of the funnel, but in the current trials, no fiber could be found stuck to the walls of the printed funnel via inspection, and so the unprocessed printed funnel was deemed sufficient. Additionally, it is important to ensure that the funnel is sealed well where it meets the scanner.

Additionally, it is important to ensure that the funnel is sealed well where it meets the scanner. Failure to do so will result in air traveling out the gaps, resulting in a higher concentration of fiber near the gaps.Figure 4: CAD drawing of the fiber dispersion funnel5. Image captureImaging was performed on a Canon CanoScan LiDE 400 at 2400 dpi. The plastic back plate was painted a matte black to provide a higher contrast for the fiber. Huntz brand A4 sized, 4 mm thick, self-adhesive plastic laminating sheets were laid sticky-side up on the scanner before fiber dispersion. This held fibers in place where they fell, ensuring none could be scattered by air in the lab, and allowed for the rapid swapping of samples. Glass panes totaling about 1/4 inch thick were placed on top of the fibers to sandwich them in place against the scanner bed, and to bring the background out of focus, which may have small scratches in the paint that could be mistaken as fiber.

In addition, the thickness of the glass panes further separated the scanner surface from the back plate, which seemed to enhance the visibility of the fibers, perhaps by allowing the glass filament to reflect the light from the scanner bar . Similar results were seen by scanning with the back plate open and a dark box covering the scanner. An example scanned image can be seen in Figure 5.261Figure 5: Scanned image example with enhanced contrast for clarity. The duller fibers are out of plane to some degree; probably stacked upon the brighter fibers.6. Fiber length measurementThe scanned images were processed and hand-measured using ImageJ [18]. To ensure the measured distribution was statistically representative, a minimum of 2000 fibers were measured per sample, as specified in ISO:22314, although the distribution seemed to begin to stabilize at around 500 fibers [19].Ideally, the entire image would need to be measured to obtain a truly representative distribution.

However, this is not practical to do by hand, as a single epoxy plug can potentially contain several tens of thousands of fibers. To reduce bias induced by measuring a smaller number of fibers, images are divided into grided subsections by multiples of the original fiber length. 45 times the square of the original length (4 mm) was chosen as the grid size (720 mm2), which contained on average 600 – 800 fiber per grid. Assuming that the spatial distribution of all fibers was uniform due to the funnel system used in step 4, any gridded subsect should contain a representative length distribution. To remove potential bias induced by sampling fibers intersecting the grid lines, it was decided that overlaps on the top and left sides would not be measured and overlaps on the bottom and right sides would be measured as shown in Figure 6.

This removes the potential to add bias by measuring more long fiber, which are more likely to be measured if crossing a border, and short fiber which are more likely to not cross a border. An added benefit of measuring border overlaps this way is that each border can correspond to a unique grid, so neighboring grid squares can be measured without interfering with shared fibers. In this study, grids squares were measured diagonally from the centermost square to the bottom right of the image.262Figure 6: Grided image example showing the border measurement classifications7. AnalysisTwo statistics were used to represent the fiber length distributions: the arithmetic mean and the length-weighted mean.The arithmetic mean is a representation of the number of shorter fibers, calculated by the sum of the individual fibers divided by the total number of fibers measured.

The length-weighted mean is calculated as the sum of the individual fiber lengths squared divided by the sum of the individual fiber lengths. The arithmetic mean is an indicator of the presence of short fibers, and the length-weighted mean gives more importance to longer fibers. Both values are often included in fiber length analyses and are useful in comparing the degree of length reduction of fiber on both sides of the distribution.𝑳𝒏 =∑ 𝑵𝒊𝒍𝒊 ∑ 𝑵𝒊𝐿𝑤 =∑ 𝑁𝑖𝑙𝑖 ∑ 𝑁𝑖𝑙𝑖2263(1)(2)Results and DiscussionFiber Length Distribution vs. Screw SpeedA summary of the results of the fiber length analysis can be seen in Figure 7 and Table 1. The length data was first validated by measuring the FLD of the 4 mm pultruded pellets. The results show that 62% of the observed fiber are within 0.2 mm of 4 mm, and 29% are 0.2 mm within 0.3 mm. The first condition tested was the 40 RPM screw speed.

The results show that 62% of the observed fiber are within 0.2 mm of 4 mm, and 29% are 0.2 mm within 0.3 mm. The first condition tested was the 40 RPM screw speed. Considering that the initial fiber length began at 4 mm, the results show that the process of extruding the material is torturous to the fiber. The 40 RPM condition exhibits a bimodal distribution with a high concentration of fibers near the original length and 0.4 mm. Certain theories on fiber breakage mechanisms may explain the large disparity. For example, Mittal et al. and Yilmazer et al. suggest that given the high fiber density of the material used, fiber bundles may survive and protect each other from bending and breaking during extrusion [8,20]. The 75 RPM samples have a similar distribution to the 40 RPM, likely due to the relatively small increase in screw speed.

The 75 RPM samples have a similar distribution to the 40 RPM, likely due to the relatively small increase in screw speed. The 150 RPM results show that the longer fiber mode experienced a significant reduction and had the largest body of short fibers of all the samples near the 0.4 mm length. The increase in shear from the 75 RPM condition seems to have overcome any mechanisms preserving the longer fibers in the 40 RPM and 75 RPM conditions.Table 1: Fiber length average representations from the feedstock and tested screw speedsFigure 7: Comparison chart of the fiber length distributions for the 40, 75, and 150 RPM conditions.264Tensile Properties vs. Screw SpeedFigure 8 exhibits the relationship between screw speed on the elastic moduli and ultimate tensile strengths of the tensile specimens respectively. A trend line of the length weighted fiber average is added to the plots to compare the relationship between the average fiber length and tensile properties.

A trend line of the length weighted fiber average is added to the plots to compare the relationship between the average fiber length and tensile properties. As expected, the results show that mechanical performance decreases with shorter fiber length. For the meager difference in fiber length between the 40 RPM and 75 RPM conditions, the tensile properties are roughly equivalent to each other. In the 150 RPM condition, doubling the RPM from 75 resulted in a decrease of about 0.5 mm in both metrics for average fiber length, and a reduction in modulus and strength by about half. It is interesting to note that 150 RPM is near the upper limit of what is a reasonable speed for the BAAM system.

Screw speed was found to have a meaningful effect on the fiber length and mechanical properties of fiber reinforced large-format printed material under the conditions of this experiment. It was seen that lower screw speeds did not significantly affect fiber attrition; however, there is an inflection point at which the influence of screw speed increases. Further investigations that vary initial fiber length and screw speeds will help to identify whether there is a fiber length retention zone in which screw speed does not significantly affect fiber attrition and how it relates to the inflection point. This will provide an analog for the degree of fiber length attrition common in large-format additive manufacturing for the GF-TPU material used here and in similar composite feedstocks.

This will provide an analog for the degree of fiber length attrition common in large-format additive manufacturing for the GF-TPU material used here and in similar composite feedstocks. Further work aims to implement deep learning to create a fiber length analysis network capable of measuring the entirety of an epoxy plug’s FLD, cementing the validity of the fiber measurement methods used in this experiment.265AcknowledgementsThis research was sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05- 00OR22725 with UT-Battelle, LLC. This research was also supported by the National Science Foundation under Award No. 2055529.References [1] L.J. Love, V. Kunc, O. Rios, C.E. Duty, A.M. Elliott, B.K. Post, R.J. Smith, C.A. Blue, The importance of carbon fiber to polymer additive manufacturing, J. Mater. Res. 29 (2014) 1893–1898. https://doi.org/10.1557/jmr.2014.212.[2] A. Durin, P.

However, information regarding flow additives and their load is limited, as is an understanding of their impact on processing conditions. This study investigates the flow behavior using static and dynamic measurements under process conditions, focusing on the influence of flow additives. Subsequently, processing studies are conducted using thermography to analyze the laser-material interaction. The characteristics of parts produced from Polypropylene and Polyamide 12 systems are also examined.

The characteristics of parts produced from Polypropylene and Polyamide 12 systems are also examined. The findings of this research enhance the understanding of the impact of flow additives on the processing conditions of laser-based powder bed fusion of polymers, potentially leading to optimized process parameters and improved part quality and mechanical properties.IntroductionThe term “Additive Manufacturing" (AM) refers to a wide range of technologies that all share a unique building method; it is a form of manufacturing where the three-dimensional (3D) product is often built into its designed shape using a layer-by-layer approach directly from a computer-aided design (CAD) model.

This makes AM fundamentally different from conventional formative or subtractive methods because it is a tool-free process, and it has several independent units creating a significantly more significant degree of complexity in terms of cost-effectiveness [1].The Laser-Based Powder Bed Fusion of Polymers (PBF-LB/P) method is widely recognized as one of the most general additive manufacturing processes employed for the fabrication of functional components, thanks to its favorable mechanical properties [2]. The polymer powder feedstock material is deposited onto a heated build platform using a powder spreading unit, such as a blade or a roller. After the powder has reached the defined build chamber temperature, exposure of the desired geometry is carried out by a laser, typically a CO2 laser. The preheated powder melts and fuses together, resulting in a dense melt, while the non-irradiated powder acts as a support structure. This process is repeated until the build job is complete.

The preheated powder melts and fuses together, resulting in a dense melt, while the non-irradiated powder acts as a support structure. This process is repeated until the build job is complete. The boundary conditions of the PBF-LB/P process necessitate specific feedstock material characteristics. In the context of PBF-LB/P, polymer properties have been categorized as intrinsic or extrinsic [3, 4]. Intrinsic characteristics include thermal, rheological, and optical properties. Extrinsic powder properties play a crucial role in achieving a homogeneous powder bed after powder application [5].State of the ArtEach laser-based powder bed fusion cycle begins by applying a powder layer to the build area. Consequently, the powder properties have a key role in the process [6].

Consequently, the powder properties have a key role in the process [6]. Several studies conclude that a homogeneous powder bed before fusion is necessary to reduce the void spaces between the particles filled during12coalescence to achieve a fully dense part with desired mechanical properties and sufficient surface quality [7, 8]. The density of the powder layer is closely related to the density of produced parts, as shown by Schmid et al. [9]. The flow behavior is influenced by various powder properties, including particle size distribution (PSD). Dechet et al. [10] compared different particle sizes of PA11 by sieving, followed by dry coating with 0.5 wt% hydrophobic fumed silica, and concluded that PSD narrowing and fumed silica coating could significantly improve the flowability of the PA11 powder. In another study by Schmidt et al.

In another study by Schmidt et al. [11], a process chain that includes wet grinding, followed by a rounding process in a downer reactor, and a dry coating step was developed for Polystyrene. The flowability of the polymer powders significantly improved after rounding and dry coating, as demonstrated by reduced tensile strength.Dadbaksh et al. [12] used fine-powder thermoplastic polyurethane (TPU) with flow additives in comparison with coarse powders and showed an increase in processing temperature and in part density. Ziegelmeier et al. [8] investigated the influence of bulk behavior and flowability on the final part properties and concluded that the surface roughness and mechanical performance of the manufactured parts depend on the powder density and surface quality.

In addition, improving the density and flowability of the powder had a positive effect on reducing porosity and pore volume.It is standard to dry coat the particles with flow additives to generate a good flow behavior which leads to a dense powder bed. Dry powder coating involves the creation of surface-functionalized composite particles by combining host particles with nanoscale guest particles. The composite particles are formed through the adhesion of nanoparticles onto the host particles, introducing roughness. This results in a considerable reduction of Van der Waals interaction forces between particles, consequently enhancing flowability [13]. It is demonstrated that there is an optimum range of flow aid concentration where flowability is maximized, and below a certain level, flowability is not significantly improved. The addition of further amounts of additives leads to a noticeable improvement, but there is a critical limit at which the positive effect diminishes [14, 15].

The addition of further amounts of additives leads to a noticeable improvement, but there is a critical limit at which the positive effect diminishes [14, 15]. In addition to that optimum range, Kleijnen et al. [16] observed a delayed coalescence of the particles via hot stage microscopy. Nevertheless, there are only a few studies that look at the influences of the entire process and show which interactions occur.In addition, there is a fundamental challenge in the existing evaluation of the flow properties of powders destined for the PBF-LP/P. While several techniques exist for measuring powder flow, the flowability of a powder is highly dependent on the nature of the applied flow field, and virtually none of the available techniques exhibit the boundary conditions of the process [6, 17]. Van den Eynde et al.

Van den Eynde et al. [17, 18] therefore reproduced an application system on a laboratory scale that simulates the application; these experiments lead to the determination of the weight of the applied layer and thus the packing density was calculated. Additionally, the influence depending on the geometry as well as the particle shape was observed. Unfortunately, there are no further studies with this setup that take into account process variables such as coating thickness, temperature, application speeds, etc.Therefore, this paper presents a new setup that mimics powder application in a PBF-LP/P machine, and provides a proper evaluation of the powder flowability for this process via surface measurements of the powder bed. Furthermore, with this equipment, different powders with different additive content are investigated, subsequently processed, and examined by means of a thermocamera to investigate influences on the process in more detail.

Finally, component properties are evaluated by means of density.13Experimental ProceduresTo investigate the effect of the different flow aids and loads on the processing and component properties, two types of additives were added separately and together to an isotactic homo Polypropylene (PP) from LyondellBasell and a Polyamid12 (PA12) from Evonik feedstock materials. Both materials were cryogenically ground, and the flow additives were added in concentrations of 0.05 wt.% and 0.1 wt.% and dry-mixed using a shaker at 49 rpm [19]. The additives utilized in PBF/LB-P have a dual purpose. Firstly, they function as distance- increasing agents to reduce particle adhesion, thereby enhancing the flowability of the powder (SiO2, D17, Evonik). This property helps to mitigate any potential clumping or agglomeration issues during the powder bed fusion/layer-by-layer processing.

This property helps to mitigate any potential clumping or agglomeration issues during the powder bed fusion/layer-by-layer processing. Additionally, the inclusion of pyrogenic alumina (Al2O3, AluC, Evonik) in the additives serves to prevent the accumulation of electromagnetic charge, which could otherwise result in inadequate powder application. By minimizing electrostatic effects, the pyrogenic alumina promotes a more controlled and uniform distribution of the powder during the printing process. Both additives are from Evonik.

By minimizing electrostatic effects, the pyrogenic alumina promotes a more controlled and uniform distribution of the powder during the printing process. Both additives are from Evonik. Table 1 displays the different composite materials for PP and PA12 used in this work.Table 1: Material combinations used in this studyMixture number 1 2 3 4 5 6 7 8 9Al2O3 in wt.% Ø Ø Ø 0.05 0.05 0.05 0.1 0.1 0.1SiO2 in wt.% Ø 0.05 0.1 Ø 0.05 0.1 Ø 0.05 0.1Particle Shape and Particle Size DistributionFor the investigation of particle size and shape of the used flow additives, scanning electron microscope (SEM) images with the Ultra Puls from Carl Zeiss AG were conducted. The images of the gold sputtered powders were recorded at a magnification of 500 and 2000 and an acceleration voltage of 5 kV.

The images of the gold sputtered powders were recorded at a magnification of 500 and 2000 and an acceleration voltage of 5 kV. The Mastersizer 2000, equipped with a Scirocco 2000 dry dispersing unit from Malvern Panalytical, was employed for laser diffraction particle sizing to determine the particle size distribution (PSD). The dispersing unit was operated at a pressure of 2 bar, and the feed rate was set to approximately 50% [20].Powder FlowabilityA reliable method for evaluating the powder flow properties is the compaction depth, as described by Hesse et al. [21]. This characteristic value, which is highly reproducible, provides valuable insights into the static properties of the powder [21, 22]. Powder flowability investigations were performed on a Discovery HR-2 rheometer (TA Instruments). The setup consists of an upper 25 mm plate parallel to a container with an inner diameter of 27 mm. The powder was sieved loose and unpacked into the container.

The setup consists of an upper 25 mm plate parallel to a container with an inner diameter of 27 mm. The powder was sieved loose and unpacked into the container. To establish a defined baseline, which results in better comparability, a normal force of 0.2 N was applied to the bulk powder. Afterward, the upper parallel plate was moved with a constant compression rate of 1 μm/s into the powder bed, and the normal force was observed. The experiments were conducted at room temperature (22 °C). For a qualitative comparison of the systems in addition to the force progression, the distance passed for a defined force difference (from 0.2 to 2.5 N) was used.Powder Application SystemAs in the state of the art, the challenge with many measurement methods is the transfer to the process, since many measurement methods are also carried out statically or under high loads, such as the shear cell, which only conditionally correspond to the boundary conditions of the process.

Therefore, this paper presents a set-up in Figure 1 that reflects the boundary conditions in the process. Ultimately, it is the construction of a PBF-LP/P14machine, but instead of the laser irradiation, it has a 3D scanner (structured light projection - fringe projection) that records the powder surface for analysis. The powder is supplied by the feeder tank, which can provide various defined quantities, followed by the recoater, which applies the powder to the free adjustable build chamber. Various application systems (roller, various blade geometries) are tested here. The speed of the application is also freely defined. In addition, the feed chamber, the build chamber and the powder surface temperature can be adjusted. For the powder surface, a pyrometer is used as in the process, and the irradiation is carried out by an infrared heater. Both chambers have thermocouples for temperature control. In addition, the powder chamber is also flushed with nitrogen.

Both chambers have thermocouples for temperature control. In addition, the powder chamber is also flushed with nitrogen. In this study, the different materials with additive content were investigated with a blade, an application speed of 150 mm/s, and a layer thickness of 120 µmm. For this purpose, 5 repetitions were carried out at room temperature. For comparison, the different systems’ surface roughness with areas of 35 mm x 35 mm were compared with each other. The analysis was carried out with a MountainsMap from Digital Surf.Figure 1: Schematic illustration of the powder application system setupThermal InvestigationsTo determine whether the additives have an influence on the thermal properties, and in particular whether they act as nucleation nuclei to increase the crystallization temperature and accelerate the crystallization process, thermal investigations were carried out with differential scanning calorimetry (DSC) measurements.

The polymer (2−5 mg) was sealed into a DSC aluminum pan and heated from 0 to 200 °C and 220 °C for PP and PA12 respectively at 10 K/min−1. After maintaining the temperature for 5 minutes to delete the thermal history, the sample was cooled to 0 °C at rates of 10 K/min−1 and then reheated using the same procedure. The reported values for the process window were obtained from the first heating and cooling cycle. For isothermal measurements, an identical heating procedure was employed. Subsequently, the samples were cooled down to the desired isothermal holding temperatures of 136 °C and 166 °C for PP and PA12 respectively at a cooling rate of 60 K·min−1. The progress of crystallization was assessed based on the crystallization half-time, which denotes the time taken for 50% of the crystallization process to occur.ProcessingThe processing experiments were carried out on a research system described by Greiner et al. [23]. The systems building space was 350 × 350 mm2.

[23]. The systems building space was 350 × 350 mm2. The machine was equipped with a 30 W CO2 laser with a focus diameter of 0.4 mm. For the conducted experiments, an inlet (100 × 100 mm2) was used to enable builds with small powder quantity. The powder coating was executed with a blade and a recoating speed of 150 mm/s. To define the process parameters, a full factorial study at three levels for single layers with 10 mm x 10 mm was performed. Therefore, the process parameters laser power (PL), hatch distance (hs), and scan speed (vs) were changed with three factors. Thereby, energy densities ED from 8 mJ/mm² to 100 mJ/mm² were covered (see Table 2). Furthermore, for PP and PA12 the building temperatures were set to 155 °C and 172 °C, respectively.𝐸𝐷 =𝑃𝐿 𝑣𝑠 ℎ𝑠(1)15Table 2: Parameter study for single layers with according running number (1-27, light grey).

The values are the energy density in mJ/mm².vS2000 mm/s1500 mm/s1000 mm/shS P0.3 mm0.2 mm0.1 mm0.3 mm0.2 mm0.1 mm0.3 mm0.2 mm0.1 mm5 W 7.5 W 10 W1 7 138 13 172 8 1413 19 253 9 1525 38 504 10 1611 17 225 11 1717 25 336 12 1833 50 6719 25 2217 25 3320 26 2325 38 5021 27 2450 75 100The best process parameter set was defined when the target values in Minitab were set to reach the determined area of 100 mm², the ratio of length to width was 1, and the lowest porosity for each single layer was obtained. All these values were obtained by optical analyses using OpenCV python. An energy density of 33 mJ/mm² (set 17) was selected for PA12, and an energy density of 22 mJ/mm² (set 22) for PP.Thermal InvestigationsAn infrared (IR) thermographic system Velox 1310 k SM (IRCAM GmbH, Erlangen, Germany) was applied for characterizing the powder laser interaction.

The system functions within the short and medium wavelength spectrums, specifically ranging from 1.5 to 5.5 µm. Greiner et al. [23] provide a schematic depiction of the assembly configuration. To safeguard the indium antimonide (InSb) detector, a sapphire glass window was positioned in front of the IR camera. The sapphire glass is not transparent to radiation of wavelengths above 6 µm, thereby filtering out the CO2 laser wavelength (λCO2 = 10.6 µm). The frame rate was set to 150 Hz, with a pixel size of 100 x 100 µm. In all experiments, a simplifying assumption was employed to maintain a constant emission coefficient and adjust it based on the initial temperature of the build chamber; this was determined using the system's pyrometer. This approach attempts to investigate whether different packing densities due to different flow aids as well as the flow aids themselves have a direct influence on the laser material interaction.

This approach attempts to investigate whether different packing densities due to different flow aids as well as the flow aids themselves have a direct influence on the laser material interaction. Here, two factors are taken into account: the maximum temperature reached during exposure to each parameter, and the average temperature experienced throughout the exposure.Part DensityTo determine the density of the samples and gain insights into their porosity, Archimedes measurements were performed [24]. The technique relies on the principle of Archimedes, where the volume of liquid displaced is equivalent to the immersed volume of the sample. This principle enables the determination of sample volume and density. The samples were weighed both in air and in a liquid medium.16Results and DiscussionPowder ShapeFigure 2 shows SEM images of the PP pure and dry blended with flow aids.

The samples were weighed both in air and in a liquid medium.16Results and DiscussionPowder ShapeFigure 2 shows SEM images of the PP pure and dry blended with flow aids. The particles display an edged surface which is explained by the powder manufacturing of cryogenic grinding. As seen with further magnification, the flow additives are homogeneously distributed on the host particle. The particle size distributions of the grinded powders are depicted as the numeric weighted cumulative sum distribution Q0(x) and the numeric density distribution weighted numeric q0(x) in Figure 2c. The particle size distribution is 83 µm and 88 µm for PA12 and PP respectively.

The particle size distribution is 83 µm and 88 µm for PA12 and PP respectively. Overall the particle size distribution of the two systems is similar with a higher fines content for the PA12.PA12sum distribtion Q3(x) in -particle size in µmPPc)11010010000.00.20.40.60.81.00.0000.0020.0040.0060.0080.0100.012density distribtion q0(x) in %Figure 2: SEM images of the grinded PP, without flow aids (a) with flow aids 0.1 wt.% Al2O3 and SiO2 (b), and the numeric weighed sum distribution and numeric density distribution of the PP and PA12 powder (c)Powder FlowabilityCompression DepthsFigure 3a illustrates schematically the development of normal force during uniaxial loading in the powder bed, comparing pure PP with PP containing flow additives. The relative change in height is measured based on the depth of the indenter as it applies a constant compression rate of 1 μm/s to a loosely consolidated powder bed with an initial normal force of 0.2 N.

When examining the change in height, it is observed that the unadded PP requires slightly increased normal forces. On the other hand, the PP with additives necessitates higher forces for the same height changes. This disparity is attributed to the fact that the powder coated with flow additives possesses a greater powder bulk density compared to the pure powder. Consequently, a greater normal force is required to achieve the same level of height change.

Consequently, a greater normal force is required to achieve the same level of height change. For an assessment of the influence of the different systems as well as possible interactions, an overview is given in Figure 3b and c, where the compression depth is given for the force differentials from 0.2 N to 2.5 N.PP + additivesa)relative height change in µm0501001502000.00.51.01.52.02.5PPnormal force in N0.1relative height change in µmSiO2 in wt.%b)PPAl2O3 in wt.%0.000.050.10801001201401601802000.000.050.1SiO2 in wt.%Al2O3 in wt.%PA12c)relative height change in µm0.050.000.000.050.10140160180200220240260280Figure 3: Schematic illustration of rise of normal force during a constant compression rate of the pure PP and dry-coated PP (a), relative height change for comparison for PP (b) and PA12 (c) with n = 517If the main effects are considered independently of each other, the SiO2 has a greater influence on PP.

However, the Al2O3 reduces not only the electric charge but also the Van der Waals forces of the particles leading to a reduced compression depth. Moreover, it is found that a stepwise increase from 0.05 to 0.1 wt.% has only a small effect, indicating that further additions beyond this point have no significant effect, this is especially true for PP. While observing the interaction between the two additives, a comparable behavior is seen, with a minimum height change reached at 0.05 Al2O3 together with 0.05 wt.% and 0.1 wt.% SiO2, with no further improvement as the Al2O3 content is increased.In the case of PA12, the impact of Al2O3 is particularly intriguing as it surpasses the effect on compressibility observed in PP and holds a stronger influence than SiO2. However, when the concentration is further increased to 0.1 wt.%, the improvement becomes less pronounced.

However, when the concentration is further increased to 0.1 wt.%, the improvement becomes less pronounced. When examining the combination of both additives, it is evident that there is no significant enhancement when transitioning from 0.05 wt.% of either component to a higher concentration blend, regardless of the proportions.Powder Flowability There is a noticeable difference amongst all compounds when comparing the compression depth of PP and PA12. This finding is supported by the application system's results, which consistently demonstrate that PP has a significantly lower surface roughness than PA12. This discrepancy implies a higher degree of homogeneity in the surface characteristics of the powder bed. It is evident that increasing the Al2O3 content alone to 0.1 wt.% is not an improvement, even when considering the interaction with SiO2 blends.

It is evident that increasing the Al2O3 content alone to 0.1 wt.% is not an improvement, even when considering the interaction with SiO2 blends. For PP, this leads to an optimum coating of 0.1 SiO2 and 0.5 wt.% Al2O3.If the additives are first examined separately for PA12, comparable effects like the compaction depth are also noticed here, where only a marginal improvement with further increase is detected.

Even more intensively than for the compaction depth, it is shown that a further increase of SiO2 in combination with Al2O3 leads to a deterioration of the powder surface.0.000.050.10304050600.1surface roughness Sa in µmSiO2 in wt.%a)Al2O3 in wt.%PP0.050.000.000.050.104050607080900.1surface roughness Sa in µmSiO2 in wt.%PA12Al2O3 in wt.%b)0.050.00Figure 4: Surface roughness depending of different additive content for PP (a) and PA12 (b) with powder flowability set up (n = 5)Thermal InvestigationsThe thermal characteristics of a polymer play a crucial role in determining the process parameters and overall process suitability. Throughout this temperature range, the molten material and the surrounding solid powder coexist. To ensure an adequate process window, the crystallization temperature of the polymer must be significantly lower than its melting temperature. This difference, known as the onset temperatures of crystallization and melting, is crucial.

This difference, known as the onset temperatures of crystallization and melting, is crucial. Ideally, polymers with a wide process window are preferred as they are less sensitive to temperature variations during the process, resulting in increased process robustness. Therefore, it is essential that additives do not act as nuclei and decrease this process window.Figure 5a-c illustrates the dynamic thermogram, which captures the first heating and cooling cycle as well as the isothermal crystallization periods of the powder for 0.1 wt.% Al2O3 and SiO2 respectively. When subjected to a heating and cooling rate of 10 K/min, the dynamic measurement revealed melting peaks at 178 °C / 164 °C, and a crystallization point at approximately 154 °C / 119 °C for PA12 and PP systems. Although not all measurements are illustrated in the graph for readability purposes, the results were seen for all possible18combinations.

Although not all measurements are illustrated in the graph for readability purposes, the results were seen for all possible18combinations. No influence of the additives compared to the pure material is detected with the measurement program carried out, but it is important to note that the selected cooling rate is considered overly high for the process. During the generation of layered components, it is important to maintain a quasi-isothermal process for a certain time period between the melting and crystallization temperatures.Therefore, Figure 5c shows isothermal measurements that are more representative of the process and also more sensitive to changes. In general, the same statement applies for the dynamic measurements, and no significant difference is seen. Considering the typical partial crystallization time t1/2, there is almost no difference between the high additive powders and the uncoated powder (having less than a 30 s difference).

Considering the typical partial crystallization time t1/2, there is almost no difference between the high additive powders and the uncoated powder (having less than a 30 s difference). It is concluded that these two additives do not show any nucleation effect for the PP and the PA12.coolingPP + 0.1 wt.% Al2O3exo upa)heat flow in W/gtemperature in °C75100125150175-2-101234PP + 0.1 wt.% SiO2heatingPPtemperature in °CPA12coolingb)exo upheatingPA12 + 0.1 wt.% Al2O3100125150175200-2-101234heat flow in W/gPA12 + 0.1 wt.% SiO2Figure 5: Dynamic DSC measurements with 10 K/min for PP with and without flow additives (a) and PA12 (b), isothermal crystallization at 166 °C for PA12 and 134 °C for PP (c)Infrared InvestigationsFigure 6a shows the thermal image after exposure for the material combinations of PA12. More details are given in Figures 6b and 6c, which show the maximum temperature and the mean temperature material peak for each material.

More details are given in Figures 6b and 6c, which show the maximum temperature and the mean temperature material peak for each material. The differences in the measured temperatures are dependent on the material systems, and amount to only a few °C which is not classified as relevant. A variation of the parameters at the same energy density leads to a much more pronounced change compared to the additive influence [25].These results and findings are also observed for PP, hence why the influence of the additives on themeasured temperature is considered negligible.a)0.10peak temperature in °CSiO2 in wt.%Al2O3 in wt.%max. temperatureb)0.000.000.050.102152162172182192202212222232242250.050.00peak temperature in °CSiO2 in wt.%Al2O3 in wt.%0.10c)mean temperature0.050.000.050.10195196197198199200Figure 6: Thermal investigations of the powder bed surface in dependence of the different coating contents, Infrared image (a), max.

temperature reached (b), and mean temperature (c) for PA1219Component DensityFigure 7 shows the measured component densities of the two material systems and their additive combinations with the same processing parameters for PP and PA12. The dependence of component densities in both PP and PA12 on the additive content of the starting powder is evident. In the case of PP, there is an increase in component density from 0.86 to 0.88 g/cm³. PA12 also exhibits an increase in component density from 0.98 to 0.99 g/cm³. Similarly to the flow analysis, it is seen that for both materials, a significant increase in the individual flow additives yields only a minor increase in density.For both PP and PA12, it is observed that increasing SiO2 to 0.05 leads to a rise in density, but it has little effect when increased past the stated range. Additionally, a reduction in the standard deviation is observed for the powders with additives.

Additionally, a reduction in the standard deviation is observed for the powders with additives. The observed differences in component density are attributed to a decrease in bulk density and an associated increase in the initial pore diameter.0.10density in g/cm³SiO2 in wt.%PPa)0.000.050.100.850.860.870.880.890.900.910.92Al2O3 in wt.%0.050.000.050.10density in g/cm³SiO2 in wt.%b)PA120.000.000.050.100.970.980.991.001.011.02Al2O3 in wt.%Figure 7: Density via Archimedes measurements for PP (a) and PA12 (b)Summary and OutlookIn this study, the impact of different additives and loads on PP and PA12 were investigated for flowability, thermal properties, laser material interaction, and part density. Flowability studies were conducted with compression depths to reveal that a further increasing of additives does not lead to significant improvements.

Flowability studies were conducted with compression depths to reveal that a further increasing of additives does not lead to significant improvements. To compare these lab values, a new experimental set up was introduced which is able to mimic the process as close as possible. These measurements showed similar trends. It is shown that the Al2O3 improves the bulk density, and that best results are gained when both systems are added. Thermal investigations by means of dynamic and isothermal measurements showed no significant impact of the chosen systems to determine that a recoating of used powder with additives is possible since no earlier crystallization is expected. During processing, the laser material interactions was observed by Infrared thermography to investigate the different packing densities due to different flow aids and to study the flow aids themselves and if they have a direct influence on the process. The results indicate a neglect able impact on the temperatures.

The results indicate a neglect able impact on the temperatures. As expected the part density showed a strong dependence of the additives. However, like the flowability studies, further increasing of the share does not lead to denser part. Additional important influences to examine subsequently are the surface roughness and the mechanics of the components. To study the transfer of flow properties in more detail, it is necessary to increase the temperature in subsequent studies, and to take into account other influences such as velocity or layer thickness.AcknowledgementThis study is funded within the IGF project (21598 N) with the title "Improving the aging stability of plastic powders the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation) – Project-ID 61375930 – SFB 814, sub-projectfor selectivelaser sintering bytailor-made additive systems" and by20T10.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAnalysis of hybrid manufacturing systems based on additive manufacturing technologyMarlon Wesley Machado Cunico, Ph.D. , marloncunico@yahoo.com.br Universidade de São Paulo campus São Carlos Concep3DDalton Alexandre Kai, M.Sc. Pontifical University Catholic of ParanaAbstract: Along the last year, additive manufacturing technologies has been proving to be a real game changer in several market segments. Nevertheless, the main foundation of production and flexible manufacturing systems generally considers classical technologies. For that reason, the present work aim to propose and investigate manufacturing systems which includes additive manufacturing technologies as part of the main or secondary production flow.

For that reason, the present work aim to propose and investigate manufacturing systems which includes additive manufacturing technologies as part of the main or secondary production flow. As result, it was identified that several marketing segments, types of components and different annual volumes tend to be better attended by hybrid flexible manufacturing systems which includes additive manufacturing technologies.Key words: Additive manufacturing, Flexible manufacturing Systems, Production1. INTRODUCTIONAlong the last years, 3d printing technologies and additive manufacturing technologies have been playing a new role in the market in addition to generate relevant changes in different market segments (WELLER et al., 2015; RAYNA e STRIUKOVA, 2016).

In spite of that, theses technologies has majorly been handle as single machine cells because of extensive fabrication time in comparison with conventional manufacturing (JAIN e KUTHE, 2013; BAUMERS et al., 2016; RAYNA e STRIUKOVA, 2016).For that reason, this work main goal is to analyse an hybrid manufacturing system that is based on additive manufacturing technologies in order to compare with conventional manufacturing systems.The Flexible Manufacturing System (FMS) can not the be considered a single technology, but a set of integrated and computer controlled technologies. According SLACK et al. (2009),the versatile of FMS provides a “capabilities package” of the system that involves the production of different products, without sequencing, without delay in the process configurations, volume production adjusted according to the demand. PANDEY et al.

PANDEY et al. (2016) features FMS in machine flexibility, production flexibility, mix flexibility, product flexibility, routing flexibility, volume flexibility and expansion flexibility.Technologies, including 3d printing and additive manufacturing technologies, allied to an FMS, results in meeting with balance a sustainable development model in the economic, environmental and social aspects that have been seeking since the end of the last century. As result, (RAYNA et al., 2015; RAYNA e STRIUKOVA, 2016) exposes that additive manufacturing tends to generate new market segments and business models.For this study, we analysed 2 cases in order to identify the effect of additive manufacturing on the total productivity metrics.

In the first case, a conventional sand casting manufacturing system is exposed while the second case includes additive manufacturing technologies into the process flow.1918In this work, we have also indicated where this hybrid manufacturing system is placed involume-variety diagram, exposing the potential market segments that might be favoured by.2. MATERIAL AND METHODSIn order to analyse the pros and cons between both of manufacturing systems, we defined abasic manufacturing layout in addition to the essential operation cost and time.In Figure 1, the basic layout of sand casting manufacturing system is presented, where the main material and product flow is also highlighted.

In this figure, patterns and core moulds are fabricated outside of plant by conventional methods generally in wood, aluminium and steel.In this study, we considered this layout as the reference scenario (case 1), while the scenario that includes additive manufacturing to fabricate patterns inside manufacturing plant is the second study case. The application of additive manufacturing for patterns, cores has already been studied in several works, where the technical feasibility was evidenced (ROSOCHOWSKI e MATUSZAK, 2000; RAO et al., 2003; CHEAH et al., 2005; MALAGI et al., 2014; RAYNA e STRIUKOVA, 2016).Figure 1 –Simple layout of sand casting manufacturing systemThe difference between those study cases is presented in Figure 2, where the general workflow of job shop sand casting manufacturing system is exposed.

Therefore, this study considers that this additive manufacturing application, which can also be considered rapid tooling, is going to be incorporated into manufacturing system .1919Figure 2 –Workflow of (a) conventional sand casting manufacturing system and (b) (a) hybrid sand casting manufacturing systemIn order to compare both cases, we assumed that all processes remain constants apart from pattern and core making. In this situation, the variation of cost and time was majorly related to tooling fabrication and life-cycle.With regards to tooling life-cycle, previous studies indicated the estimated number of parts that each sort of tooling support. For that, we analysed different types and shapes of patterns and compared against conventional and additive manufacturing based systems. In this case, patterns sizes varied from 100x100x50mm until 300x300x100mm.

In this case, patterns sizes varied from 100x100x50mm until 300x300x100mm. In Table 1, the average tool investment is presented besides average life-cycle.Table 1 - Comparison between Tooling material and maximum supported number of parts Metal based on (ROSOCHOWSKI e MATUSZAK, 2000; SOARES, 2000)Although some of values are extracted from previous works, further investigation is still neededto better understand the cost-life correlation.1920As the tooling cost is considered a capital investment, the contribution of tooling for the part cost is amortised by the volume of parts which is fabricated. Therefore, the total tooling cost might ) and annual be defined as a function of moulding tool cost ( parts amount (N), as shown in (1) (REES, 1996; CHATAIN e DOBRACZYNSKI, 1997; FAGADE e KAZMER, 2000).

In Addition, we included the cost of Additive manufacturing equipment (Am).toolmouldingC_df), depreciation factor (.dtoolmouldingfNCC_Therefore, the amortized product cost by product can be found by Eq. (2), where n is the numberof products produced in the manufacturing system.nidntoolmouldingnnfNCC1_3. RESULTS AND DISCUSSIONSIn this study, we identified that although tooling that was fabricated by additive manufacturing technologies implies on reduced life, total cost was found to be lower than conventional tooling making. This can be seen in Figure 3, where a comparison between cost amortization against the number of fabricated parts is presented.Figure 3 –Comparison between tooling fabricated by conventional and additive manufacturing technologiesOn the other hand, Figure 5 compares the cost of 1 product conventional pattern making against 1, 2 and 5 products additive manufacturing pattern making.

In this case, it is possible to see that amortized cost of 5 products tooling that were fabricated by additive manufacturing is still lower than conventional cost designated for only 1 product. In contrast, this behaviour can be observed for small and medium product volumes.1921(1)(2)Figure 4 –Comparison between tooling fabricated by conventional and additive manufacturing technologies in accordance with the product varietyAs result of this first analysis, we can identify a new manufacturing profile that implies in the constant tooling making. Therefore, each moment that a new tool is needed to be fabricated generates new opportunity to change part or product and increase variety.

Therefore, each moment that a new tool is needed to be fabricated generates new opportunity to change part or product and increase variety. The main implication of this can be seen in Figure 5, where a comparison between conventional and hybrid manufacturing based in additive manufacturing is exposed.Therefore, additive manufacturing was proof to attend either small and medium volumes withhigh variety or high volumes with low variety.Figure 5 – Analysis of Product variety over the time against production volumeAs consequence of this new approach, new economical models might be introduced and new opportunities can be exploited. At the same way, this new manufacturing approach also allows the growing of customer interference into design, whereas new product updates can be easily implemented during the product life.4.

At the same way, this new manufacturing approach also allows the growing of customer interference into design, whereas new product updates can be easily implemented during the product life.4. CONCLUSIONS As consequence of this work, the comparison of conventional and hybrid manufacturing systems which includes additive manufacturing technologies as part of process was analysed.

CONCLUSIONS As consequence of this work, the comparison of conventional and hybrid manufacturing systems which includes additive manufacturing technologies as part of process was analysed. In addition,1922the economical analysis of pattern making was performed and additive manufacturing tooling was identified as more profitable.The analysis of variety was also done and additive manufacturing tooling making was found togenerate higher variety than conventional tooling with the same cost.Therefore, additive manufacturing was proof to attend either small and medium volumes withhigh variety or high volumes with low variety.Additionally, this new manufacturing approach generates a new economical model that might be a game changer, whereas it increases the collaboration of customers into design, allowing high amount product updates and product changes during the product life cycle besides ensuring conventional manufacturing strategies.5. REFERENCESBAUMERS, M., P.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAnalysis of Layer Arrangements of Aesthetic Dentures as a Basis for Introducing Additive ManufacturingK. Pitz* and R. Anderl*Department of Computer Integrated Design, Technische Universität Darmstadt, 64287 Darmstadt, GermanyAbstractAesthetic dentures are highly individual products, needing to adapt to the patient’s natural dentition in both color and shape. In general, dentures are geometrically complex structures, consisting of several layers of different materials, which are applied consecutively. This is why efforts are being made to introduce additive manufacturing into the production process of aesthetic dentures which is to date largely based on manual work of skilled professionals. In order to successfully apply additive manufacturing accurate models of the denture to be produced are essential.

In order to successfully apply additive manufacturing accurate models of the denture to be produced are essential. In this paper the layer arrangements of dental crowns from real patient cases are analyzed. Different approaches, based on normal vectors, cylindrical coordinates, spherical coordinates and paraxial rays, are compared to each other. The emphasis of this comparison lies on transferring layering strategies from analyzed patient cases to future patient cases where the geometry might be slightly different but a similar optical impression is desired. The most suitable approach for this purpose has been identified, implemented and tested.Keywords: Dentures, CAD/CAM, Layering strategies, Thickness measurementsIntroductionThe potential for additive manufacturing in dentistry is immense. Dentures are of highly individual shape and are produced in quantities of one [7; 9; 10].

Dentures are of highly individual shape and are produced in quantities of one [7; 9; 10]. Layering strategies must be applied to mimic natural properties of teeth and allow for a wide variety of shades [22; 23]. The strengths of additive manufacturing – capabilities for complex and small geometries [5], method- immanent use of layers [19], energy and material efficiency [25], wide range of usable materials and material combinations [5] – can be optimally exploited. Production time and costs can potentially be reduced as well as the quality of the final product improved if a shift is being made from manual work of dental technicians to machine-based production.Typical materials for dental restorations include metals, polymers and ceramics [4; 6]. This work focuses on dental crowns, i.e. replacement of single teeth, in the form of milled frameworks which are veneered with ceramic materials. Translucency is a distinctive differentiation criterion for veneering materials [1; 28].

replacement of single teeth, in the form of milled frameworks which are veneered with ceramic materials. Translucency is a distinctive differentiation criterion for veneering materials [1; 28]. The framework material is usually opaque. Materials of following layers allow for light trespassing and therefore offer the technician possibilities to create an optical impression as natural as possible. For example, lighter colors at the cutting edge of a tooth, white spots or translucent effects of yellowish and greyish colors can be realized. Apart from their translucency properties the base colors of the materials are of1079importance [20]. Professional suppliers of dental materials offer wide varieties of color palettes [17].Layering strategies are essential to create dentures of high aesthetic value. Thicknesses of layers [24] contribute to the final color impression [16]. Generally speaking, the thicker a layer is, the less visible are color and shape of the underlying layers.

Thicknesses of layers [24] contribute to the final color impression [16]. Generally speaking, the thicker a layer is, the less visible are color and shape of the underlying layers. Also, the less translucent a material is, the less visible are color and shape of the underneath. In addition, the base colors of the materials of all different layers have an influence on the color that is perceived on the outer surface of the denture. Until now, experience and skills of dental technicians are relied upon to adjust veneering layers of dentures correctly [11; 18; 29]. In order to enable the use of additive manufacturing it is important to create a digital representation of these layers. In this way, they become reproducible and can be passed on to material-applying machines. That is where this work sets in. Layer arrangements of existing dental crowns are analyzed by different approaches.

That is where this work sets in. Layer arrangements of existing dental crowns are analyzed by different approaches. The underlying aim is to be able to store layer arrangements and transfer them to future patient cases where the specific color impression is needed again.Data baseFor this work single crown dentures from 30 anonymized patient cases have been analyzed. They are all dentures to replace front teeth. These dentures have been produced by one and the same dental technician in a laboratory of current standards, namely Form For Function GmbH, located in Darmstadt, Germany. The dental technician at work digitized the dentures after every applied material. They are each available as 3D scans and photos. The files stemming from the 3D scans are in STL format. Fig. 1 gives an example of such a digitized production step.(a)(b)Fig. 1: (a) STL model and (b) photo of a dentureThe dentures analyzed in this work are based on materials from VITA Zahnfabrik GmbH.

1: (a) STL model and (b) photo of a dentureThe dentures analyzed in this work are based on materials from VITA Zahnfabrik GmbH. Color 3M2 has been chosen as it is a common color in dental practice. Enamic is used as the framework material. LC flow is used for the veneering steps. Fig. 2 depicts the used materials.1080i'N:i BASE N DENTINE~ ~,ENAMEL- " ' - 'ffi"'---VllAVM.lC .,,.(a)v1TAfNl',lf(b)Fig. 2: (a) VITA Tooth Guide 3D-Master to determine a patient’s tooth color, (b) VITA Enamic block for milled frameworks, (c) VITA VM LC flow for application of upper layers [26]The most simple layering strategy consists of the framework milled out of Enamic, one veneering layer of dentine LC flow (to model the overall body) and one veneering layer of enamel LC flow (to model the cutting edge). Additionally, the tooth stump is relevant [7]. It is not part of the denture, but part of the remaining tooth substance in the patient’s jaw.

Additionally, the tooth stump is relevant [7]. It is not part of the denture, but part of the remaining tooth substance in the patient’s jaw. The finalized denture will be fitted onto the tooth stump. The stump therefore defines the inner contour of the framework and is for that reason included into the 3D STL models. Fig. 3 depicts one example of a simple layer arrangement. Further layers can become necessary when there are more complex color effects to be modeled.Fig. 3: Example of a layer arrangement (white = tooth stump, red = framework, blue = dentine, yellow = enamel)Comparison of possible approachesDetermining the thickness of a layer as shown in Fig. 3 is a complex endeavor. Each layer is a three-dimensional object that varies in its thickness over the entire freeform surface, with the variations taking place in small value ranges.

3 is a complex endeavor. Each layer is a three-dimensional object that varies in its thickness over the entire freeform surface, with the variations taking place in small value ranges. In addition to measuring the layer thicknesses, the approach this paper aims for should also be able to transfer the layer arrangement to a future patient case. It is assumed that for such a potential future patient case the contours of the tooth stump and the outer silhouette are given. The first being determined by what is left of the original tooth and how the dentist grinds it down. The latter being determined by the adjacent teeth in the patient’s1081(c)jaw. Again 3D scanning devices, e.g. intra oral scanners, could be used to digitize these two limiting contours.When discussing thicknesses, it should be pointed out how the term thickness is understood.

intra oral scanners, could be used to digitize these two limiting contours.When discussing thicknesses, it should be pointed out how the term thickness is understood. Generally, thickness is conceived as a cross-sectional expansion of a body [27], measured in the direction of the surface normal from a lower surface to an upper surface. This paper complies with this definition. For all layers of translucent materials the actual thickness measured along the surface normal is what defines to which extent underlying structures are visible and contribute to the color impression. Only if this thickness is applied to a new denture in a new patient case it is ensured that the color impression at a certain point resembles the color impression of the original denture at the corresponding point. Whenever thicknesses are not measured in normal direction in the approaches presented in the following sections it will be pointed out in their description.I.

Whenever thicknesses are not measured in normal direction in the approaches presented in the following sections it will be pointed out in their description.I. Normal vector approachSince the 3D models at hand are available as STL files, i.e. with triangulated surfaces, it is an obvious choice to use the existing normal vectors of the model’s triangles. The analyzed files contain between 50.000 and 75.000 triangles per layer surface. Even higher resolutions could be achieved in the scanning process. Thus, information about layer thicknesses at an almost unlimited number of locations can the obtained. Fig. 4 shows how normal thicknesses are measured in a dental crown model.Fig. 4: Schematic representation of the normal vector approach with some exemplary measuring pointsThe main advantage of using the normal vector approach is that the correct thicknesses (i.e. thicknesses in surface normal direction) are measured and no need for error handling arises.

thicknesses in surface normal direction) are measured and no need for error handling arises. Apart from that, all areas of the tooth are included into the calculation as all normals of all triangles are used. However, one difficulty of the approach is to determine in which direction it should be applied. When working with normal vectors it does make a difference whether distances are measured from the lower to the next-highest layer or the other way around. Fig. 5 illustrates this fact.1082Fig. 5: Visualization of the directional dependency of the normal vector approach (green = thicknesses measured from the inside outwards, blue = thicknesses measured from the outside inwards)In both cases, when measuring from the inside to the outside and when measuring from the outside to the inside, extremely large values for layer thicknesses can occur. This is the case whenever a triangle exhibits a large slope. The upper arrow in Fig. 4 is an example of this phenomenon.

This is the case whenever a triangle exhibits a large slope. The upper arrow in Fig. 4 is an example of this phenomenon. These values are not to be considered incorrect as they do capture the distance to the next layer in the defined way. However, such values may cause difficulties when it comes to transferring the layer arrangement to another patient whose tooth’s inclination angles are not as severe. In even more extreme cases, the normal vector of a steep surface may not hit the next layer at all, but the current layer again. This effect occurs in particular when measuring from the outside to the inside. The respective values should be excluded because they don’t contain any information about the layer’s thickness. Fig. 6 illustrates both cases of normal vectors on steep surface areas.Fig.

The respective values should be excluded because they don’t contain any information about the layer’s thickness. Fig. 6 illustrates both cases of normal vectors on steep surface areas.Fig. 6: Examples of extreme values for the normal vector approach (orange = extreme thickness that can still be used, red = no thickness information)The task of transferring layer arrangements measured along normal vectors is problematic as a whole. The triangles of the meshes of the target denture (of the new patient) will differ greatly from those of the analyzed denture (from another patient). Both tooth stump mesh and outer contour mesh are highly individual. It cannot be assumed that similarity in position and orientation of normal vectors exists. Thus, a complex mapping of the measured thicknesses onto the new surfaces would need to be developed.1083II.

Thus, a complex mapping of the measured thicknesses onto the new surfaces would need to be developed.1083II. Cylindrical coordinates approachWorking with an approach based on cylindrical coordinates can be imagined as if an enveloping cylinder was placed around the denture to be analyzed. With the help of the coordinates z (height), ρ (radius) and φ (angle) the denture can be measured in slices. Slices increase by a defined increment Δz. A number of angles φ1 to φn has to be defined at whose positions the thicknesses of the layer arrangement are measured. The resolution of this approach can be improved by increasing the number of slices and angle positions. Fig. 7 illustrates the approach./ II____,I \ - \.../-----\t I\/ I\\fI:,. zI(a)(b)Fig.

Fig. 7 illustrates the approach./ II____,I \ - \.../-----\t I\/ I\\fI:,. zI(a)(b)Fig. 7: Schematic representation of the cylindrical coordinates approach with some exemplary measuring points; (a) side view, (b) top viewThis approach is not direction dependent as the normal vector approach is. Whether thicknesses are measured from the coordinate system center outwards or from the cylinder surface inwards makes no difference. Another advantage is that implementation becomes easier compared to the normal vector approach. No loops are needed to go through all triangular facets and gather the respective normal vectors. For this approach only the bounding cylinder has to be determined based on the dimensions of the denture. Thickness measurements are then based on this cylinder and its segmentation.A downside of the cylinder approach is that the bounding cylinder is not always easy to find.

Thickness measurements are then based on this cylinder and its segmentation.A downside of the cylinder approach is that the bounding cylinder is not always easy to find. It has to be reflected upon which position the point of origin shall take. It could either be determined purely based on the outer contour of the denture or it could be chosen to coincide with a dentistry relevant point such as positions of cusps or fissures or the tooth’s center of mass. Further, it could be argued that the orientation of the cylinder should follow the orientation of the tooth stump or that of the insertion direction of the finished denture. However, even if factors are chosen that are prominent on the tooth or significant for the dentist, the difficulty of transferring it to the next patient still arises.

However, even if factors are chosen that are prominent on the tooth or significant for the dentist, the difficulty of transferring it to the next patient still arises. As all tooth geometries are very individual, the transfer process might not yield satisfying results for the new patient.Assuming that a reference system for the cylinder has been found, there is still a necessity for a scaling step from the existing denture to the new patient case. This concerns the cylindrical slices. If the second tooth is larger than the first one, they need to be upscaled. Fig. 8 gives an example of such an upscaling process. Otherwise, if the second tooth is smaller, the slices need to be downscaled. The angles φ1 to φn are retained as they are.1084I/II \---... ___J - \..../_~~~ -✓--'\--~/ II'/ "I,___Fig.

The values here are not measured along surface normals and are therefore not as accurate when it comes to capturing what determines the color impression. Fig. 10 shows how these two types of measured thicknesses can differ from each other.Fig. 10: Deviations between normal thicknesses (green) and horizontal thicknesses (red)1085The differences become particularly large in areas with steep gradients. However, it can be argued that the error that is committed by not using normal directions for measuring can be minimized by increasing the resolution.III. Spherical coordinates approachThe approach based on spherical coordinates is similar to the one based on cylindrical coordinates. Except that an enveloping sphere is chosen here. The sphere and positions within the sphere can be described by r (radius), θ (polar angle) and Φ (azimuthal angle). The radius is defined by the size of the tooth.

The sphere and positions within the sphere can be described by r (radius), θ (polar angle) and Φ (azimuthal angle). The radius is defined by the size of the tooth. The more rays, each described by the respective pair of θi and Φi , are used, the higher the resolution. Fig. 11 shows some rays cast through a denture in a bounding sphere.Fig. 11: Schematic representation of the speherical coordinates approach with some exemplary measuring pointsThe benefit when compared to the cylindrical approach is that it does not leave regions of the tooth uncovered. Again it is independent of the direction of application. Whether thicknesses are measured from the sphere’s surface inwards or from the center point outwards is irrelevant. Implementation is as convenient as described for the cylindrical approach.

Whether thicknesses are measured from the sphere’s surface inwards or from the center point outwards is irrelevant. Implementation is as convenient as described for the cylindrical approach. The effort to create loops over all normal vectors can be avoided here as well, since only the sphere and its division are needed as a basis.However, the problem of finding an adequate reference system for transferring layer arrangements from one patient to another remains. It is as critical to find a suitable center point for the sphere as it is for the cylinder. Again the center of mass of the stump or the final denture might suggest themselves, yet they vary greatly from person to person. The only advantage gained is that the sphere only needs a center point, not the additional longitudinal direction the cylinder needs.When working with spherical coordinates no scaling is needed for a new patient case.

Only the radius of the sphere will be different, but all pairs of θi and Φi stay exactly the same. Only these rays are needed to determine the layers for the new patient along them.Similar to the cylindrical approach, it has to be taken into account that the thicknesses measured are not normal distances. Fig. 12 shows how the two types of thickness values may differ from each other. In this case, they show these deviations independent of surface inclinations.1086Fig. 12: Deviations between normal thicknesses (green) and spherical thicknesses (red)IV. Paraxial ray approachThe enveloping body for the paraxial ray approach is a cuboid. This bounding box has its dimensions defined according to the axes of a Cartesian coordinate system, typically denoted by x (length), y (width), z (height).

This bounding box has its dimensions defined according to the axes of a Cartesian coordinate system, typically denoted by x (length), y (width), z (height). The fact that Cartesian coordinates are the ones most commonly used and that today’s CAD systems offer built-in functions to create bounding boxes for a body makes this approach especially convenient to apply.Paraxial rays in this context mean straight lines that go through the bounding box. They are placed through the cuboid in all three coordinate directions, with defined intervals Δx, Δy and Δz. By reducing the intervals, the resolution of the approach can be scaled up. If there is an interest in one particular direction, e.g. because the tooth exhibits peculiarities on the cutting edge, the interval for this direction can be reduced in a targeted way. For each ray the thicknesses of the layers are measured in the direction of the ray. Fig.

For each ray the thicknesses of the layers are measured in the direction of the ray. Fig. 13 shows an example of a bounding box and the casted paraxial rays.I/----\(a)7-----~I\._,/(b)!J.zI/II/"I!J.y----"'- I'-.. [\ \ I\ I\ \ IJ I J V I /✓--- I/ ~ / I/,._ ~ \I/"-- I"'- '- '---- -- t / v....- v\ \'---~/J.x/J.xFig. 13: Schematic representation of the paraxial ray approach with some exemplary measuring points; (a) side view, (b) front viewThe major advantage of this approach is that no reference system has to be chosen if the layer arrangement is to be transferred. Only the outer contour of the existing tooth and the outer contour of the new tooth need to be known as they are what defines the bounding boxes. These contours are known by definition of the use case discussed in this paper.

These contours are known by definition of the use case discussed in this paper. However, scaling becomes necessary to determine the layers for the new patient case as it cannot be assumed that the bounding boxes will be of identical size. This is carried out by adjusting the intervals. The number of rays1087per direction stays the same so that no information is lost nor artificially generated. Fig. 14 illustrates the scaling process for the case that the new tooth is bigger than the original one.I I~f\I/VI/\I~I'-I+---I/I/Fig. 14: Transfer of a layer arrangement onto a new tooth for the paraxial ray approachAgain it is an approach that is independent of the direction in which it is carried out. It does not matter whether thicknesses are measured in or against the respective coordinate direction.

It does not matter whether thicknesses are measured in or against the respective coordinate direction. It should only be considered that in two directions there will be two sets of thicknesses (front – back and left – right) while in one coordinate direction there will be only one set of thicknesses (upwards to the cutting edge). This is due to the fact that the denture is open in the lower area where it is to be placed on the tooth stump in the patient’s jaw. It can be assumed that the entire tooth is well covered by taking into account all three coordinate directions. Implementation is even more feasible than described for the spherical and cylindrical approach.A challenge that remains is to account for the fact that distances are not measured along surface normal direction. Similar to the cylindrical approach (see Fig. 10), paraxial thicknesses may differ from normal thicknesses. Differences are especially significant when slopes are steep.

Similar to the cylindrical approach (see Fig. 10), paraxial thicknesses may differ from normal thicknesses. Differences are especially significant when slopes are steep. The issue can partially be resolved by keeping the resolution high.V. ConclusionBased on the discussion of the approaches in the above sections, a decision is being made to pursue the paraxial ray approach further. The advantage which particularly drives this decision is that this approach offers independence from a reference system. This is of great importance for the use case this paper deals with, namely the transfer of a layer arrangement from one patient to another.Furthermore, implementation of the paraxial approach provides a solid foundation for further development. It can be experimented with different resolutions and different patient cases for the transfer process.

It can be experimented with different resolutions and different patient cases for the transfer process. If it can be proven that results of sufficient quality are produced by this approach, the elaborate development of a reference system can justifiably be disregarded. If the opposite scenario occurs, the implementation of this approach can nevertheless act as a benchmark against which it can be measured how much other approaches improve accuracy.1088ImplementationFor programming purposes the CAD (Computer Aided Design)/CAM (Computer Aided Manufacturing) software Pictures by PC 3.8 [21] has been used. It comes with its own object- oriented programming language, which is based on BASIC (Beginner’s All-purpose Symbolic Instruction Code) [2].

It comes with its own object- oriented programming language, which is based on BASIC (Beginner’s All-purpose Symbolic Instruction Code) [2]. Thus, one and the same environment can be used to load, align and modify dental models as well as to apply self-written algorithms to them.When implementing any of the approaches mentioned in sections I to IV the questions of how to detect intersection points arises. Since all models are in STL format, i.e. are assembled of a multitude of triangles, and all approaches apply some kind of lines cast through the model, it is important to know when a line hits a triangle. The distance from an intersection point of a line and a triangle to the next intersection of the same line and a triangle of the next-higher layer is what defines the thickness in the respective approach.A ray-triangle-intersection (RTI) algorithm that is well established is the one proposed by Möller and Trumbore [15].

It stems from the field of computer graphics where it is commonly implemented for ray tracing computations [14]. In this work it is chosen because of its computational efficiency. The reason for this characteristic is that it does not require a pre- calculation of the plane in which the triangle containing the intersection point lies.The algorithm is based on barycentric coordinates [8]. This means that a transformation from the overall global coordinate system to a system based on the vertices of the triangle is made. In barycentric coordinates a point T that lies within a triangle defined by vertices V0, V1 and V2 can be described as follows [15]:T(u,v) = (1 – u – v)V0 + uV1 + vV2with u and v being the coordinates which must satisfy the conditions u ≥ 0, v ≥ 0 andu + v ≤ 1.A parametric definition of a line, i.e.

of the ray in this scenario, is the following [15]:R(t) = P + tDwith R being an arbitrary point on the line, P being the starting point which has been chosen to define the line, D being its direction vector and t the distance traveled from the starting point in the direction of D to end up in R.If now a triplet (t, u, v) is found that satisfies both equations it is the intersection pointbetween ray and triangle. For further implementation matters the reader is referred to [14].To ease the interaction with the user, a graphical user interface (GUI) is desired for the application presented in this paper. The user shall have the opportunity to enter values for the number of rays in x, y and z direction. The interval values Δx, Δy and Δz are then calculated by dividing the size of the bounding box by the number of rays + 1. Fig.

The interval values Δx, Δy and Δz are then calculated by dividing the size of the bounding box by the number of rays + 1. Fig. 15 shows how this GUI is realized in the Pictures by PC software.1089(1)(2)MlayerTransfermnumber o f rays in x-direction: 2~number o f rays in ~-direction: 25number o f rays in 1-direction: 15Fig. 15: Graphical user interface for the implemented layer transfer algorithmTo give an overview of the entire implementation for the analysis of the layer arrangement of a denture and for the transfer of this arrangement onto a new tooth with only the outer contour and the contour of the tooth stump known, a flowchart is given in Fig.

16: Flowchart of the implemented layer analysis and transfer algorithm; (a) entire process, (b) analysis subroutine, (c) transfer subroutineThe fact that the first rhombus in Fig. 16 (a) checks the number of mesh bodies originates from the considered use case (transfer of a layer arrangement from one patient to another). It is assumed that all dentures have at least 3 layers (see Fig. 4 for example). 3 layers equal 4 mesh bodies, namely the tooth stump contour, the framework outer contour, the dentine outer contour1090and the enamel outer contour. For the new denture it is assumed that only 2 of these meshes are known (see Fig. 8, right part, for example). Those two are the tooth stump contour as it is captured by an impression or 3D scan created at the dentist and the outer contour which is determined by the adjacent teeth and captured by the dental technician via CAD software.

The boundaries of the inner layers of the new denture are calculated by the algorithm.ResultsThicknesses of the different layers a denture consists of could successfully be measured by the paraxial ray approach and the respective implementation. Within the available patient cases as described under Data Base thicknesses of the core layer (first layer on the tooth stump, created out of VITA’s Enamic material) vary around a mean of 1.9 mm in their maximum values. Maxima of the dentine layer (middle layer, created out of VITA’s 3M2 LC flow dentine) vary around a mean of 0.2 mm. For the enamel layer (outermost layer, created out of VITA’s 3M2 LC flow enamel) maximum thicknesses fall around a mean of 0.5 mm.

For the enamel layer (outermost layer, created out of VITA’s 3M2 LC flow enamel) maximum thicknesses fall around a mean of 0.5 mm. These results are plausible when considering that the core layer is the most massive one, responsible for the structural integrity of the denture [13], and that the enamel layer is the one used for creating the cutting edge and therefore contributes to large values in height direction. Other layers for more complex color effects are not listed here, but can nevertheless be measured by the same described means.Fig. 17 gives an example of a layer transferred to a new denture. Green dots stand for the intersection points between rays and triangles which were recorded for the original denture and then applied onto the new one shown here. Another point cloud would appear on top of the green one to make the simple three-layer strategy complete.Fig.

Another point cloud would appear on top of the green one to make the simple three-layer strategy complete.Fig. 17: Result of a layer transferred onto a new dentureSummary and OutlookThe approaches to analyze layer arrangements of dentures that have been compared in this paper include a normal vector approach, a cylindrical coordinate approach, a spherical coordinate approach and a paraxial ray approach. Advantages and shortcomings of each of these approaches have been pointed out. The direction of thickness measurements and the choice of a reference1091system for the transfer to new patient cases have turned out to be a particular focus. Even though measurements along normal vectors are desirable as they are in line with the general definition of thickness, a decision has been made for the approach based on paraxial rays for transferability and implementation reasons. Implementation steps using the software environment Pictures by PC 3.8 have been explained.

Implementation steps using the software environment Pictures by PC 3.8 have been explained. Finally, exemplary results of layers measured in dentures stemming from the data base at hand have been presented.The work presented in this paper is of initial, fundamental character. Further steps must be taken in order to make the chosen and implemented approach fully operational for additive manufacturing of dental prostheses. These steps involve the creation of surfaces through given points. So far, as can be seen in Fig. 17, the layer arrangement when transferred to a new tooth is present in form of point clouds. Delaunay triangulation [3] or Marching Cubes [12] could be used to create surfaces out of these points and thus define the exact dentine and enamel layers for the new patient.Assuming that materials and machines are sufficiently adapted to be utilized for additive manufacturing in dentistry, the implementation described here could be used to produce prototypes of dentures.

With the availability of real physical prototypes it can be checked whether the optical impression, especially in terms of colors, is achieved as expected. For example, the so-produced prototypes could be compared to others from manual production, could be exposed to different light conditions or inserted into a patient’s jaw to assess the optical interaction with the natural teeth around. If the aesthetic aspirations are not met, it is recommended to compare the proposed approach to the normal vector one. The normal vector approach, even though more elaborate with regard of implementation and transfer, ensures the exact measurement of thicknesses. If necessary, modifications must be made to the existing approach, e.g.

If necessary, modifications must be made to the existing approach, e.g. adjustments of the resolution, or it may even become necessary to switch to the normal vector approach and find a suitable reference system for transferring it onto new teeth.AcknowledgmentsThis paper is based on work within the BMBF (German Federal Ministry of Education and Research) funded project MYTHOS. The authors thank all members of the project consortium, especially Form For Function GmbH for providing data from anonymized patient cases, Schott Systeme GmbH for providing the software Pictures by PC 3.8 and VITA Zahnfabrik H. Rauter GmbH & Co. KG for supporting the work with material know-how.References[1] Brodbelt, R. H. W., O'Brien, W. J., and Fan, P. L.: Translucency of Dental Porcelains. In: Journal of DentalResearch 59 (1), pp. 70–75.

alysis of Mesostructure Unit Cells Comprised of Octet-truss StructuresScott R. Johnston*, Marques Reed*, Hongqing V. Wang†, and David W. Rosen*The George W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332 †R&D Group, IronCAD, 700 Galleria Pkwy, Atlanta, GA 30339Abstract: A unit truss finite element analysis method allowing non-linear deformation is employed to analyze a unit cell comprised of n3 octet-truss structures for their stiffness and displacement compared to their relative density under loading. Axial, bending, shearing, and torsion effects are included in the analysis for each strut in the octet-truss structure which is then related to the mesostructure level (unit cell). The versatility of additive manufacturing allows for the fabrication of these complex unit cell truss structures which can be used as building blocks for macro-scale geometries.

The versatility of additive manufacturing allows for the fabrication of these complex unit cell truss structures which can be used as building blocks for macro-scale geometries. The finite element calculations are compared to experimental results for samples manufactured on a Stereolithography Apparatus (SLA) out of a standard resin.Introduction Lightweight compliant structures are becoming more desirable in many areas of industry due to the advancements in manufacturing methods that are now capable of fabricating their complex geometries. Components with an internal truss structure designed for specific loading applications is an example of a lightweight structure. The automotive and aerospace industries are very interested in lightweight structures because lightweight structures can improve their product’s performance.

The automotive and aerospace industries are very interested in lightweight structures because lightweight structures can improve their product’s performance. Truss structure applications have been limited in the past because of the large cost required for fabrication, or the structure was impossible to fabricate using traditional manufacturing methods. Additive manufacturing is now capable of fabricating lightweight truss structures, possibly with similar mechanical properties to conventional manufactured components with the same exterior component design.1The initial steps for introduction of truss structures in commercial applications requires mechanical analysis of the capabilities of the truss structures to determine their mechanical properties. This will allow designers to choose truss designs with knowledge of their strength, stiffness, and weight. Since truss structures are a series of struts and nodes, they behave differently from solid components.

Since truss structures are a series of struts and nodes, they behave differently from solid components. The development of fast computing systems has allowed Finite Element Analysis (FEA) to become a standard method for component design analysis, but most FEA codes usually use elements discretely in the form of beams, rods, plates, or solid elements. The nodes of the truss structures physically possess the loading constraints (tensile, compression, torsion, elongation, and buckling) of many different types of the aforementioned element types, but an element possessing all of the unique loading conditions which a truss structure node experiences does not currently exist.A unit truss approach FEA program has been developed at Georgia Tech [1] to specifically analyze complex truss structures (described in more detail in Section 2.4) which includes tensile, compressive, torsion, and buckling effects due to structural loading.

The unit truss FEA analysis is compared to experimental results for stiffness and displacement for a specified relative density of a unit cell. Each unit cell contains a different number of octet truss structures [2] possessing421the same relative density of approach are presented in Section 3.4..35.0(cid:32)(cid:85)A comparison of experimental and unit truss FEA2 Meso-structure design analysis of componentsComponents can be geometrically decomposed into smaller elements for analytical purpose which is precisely the analytical process employed by Finite Element Analysis (FEA). The same idea can be used to successively decompose a component (macro-scale) into medium sized elements (meso-scale) which is further subdivided into smaller building blocks (micro-scale).

The same idea can be used to successively decompose a component (macro-scale) into medium sized elements (meso-scale) which is further subdivided into smaller building blocks (micro-scale). This decomposition ideology is the basis for the unit cell design and analysis approach to lightweight components by using truss structures within the unit cells.2.1 Unit cell approach for component designA multi-scale design approach is employed to create an appropriate truss structure for specific component geometries. The geometric design of the component can be decomposed into mesostructure unit cells (structures possessing feature sizes between micro and macro-scales as displayed in Figure 1a) such that the mesostructures are used as the basic building blocks for the component geometry.

Each mesostructure unit cell is further decomposed into smaller truss structures (Figure 1c) where the octet-truss structure (Figure 1b) has been chosen to be the building block for each unit cell [3] because the mechanical behavior of the struts (lattice structure) can be predicted more readily than materials with random voids, such as foams. Performing this decomposition creates a transition from a macro scale (the component level) to a meso-scale (unit cell) and finally to the micro-scale (octet-truss). An example of this decomposition is displayed in Figure 1 to illustrate the multi-scale analysis of the component.0.50110.50 000.5(a)(b)(c)Figure 1. (a) Component geometry decomposed into unit cells, (b) octet-truss structure, and (c) a unit cell containing a 2x2x2 octet-truss structure.There are many reasons for employing the unit cell design methodology.

First, by using unit cells with known mechanical behavior, the unit cells can be used as an element type in a Finite Element Analysis (FEA) which reduces the computational intensity of the FEA calculations compared to the current FEA definition of numerous nodes and struts comprising the truss structure. A second advantage in using this mesostructure approach provides the designer local control of the mechanical behavior at the component level by altering the composition of the unit cells at the mesostructure level.

This advantage also allows for a maximum component performance with respect to weight under defined loading conditions because less dense unit cells may be used in locations that carry less of the applied load.The following investigation of unit cells comprised of arrays of octet-truss structures will expand upon previous research of the octet-truss structure by: (1) including bending, shearing, and torsion effects in addition to the tensile effects and (2) analysis of arrays of octet-truss422structures that include the interaction between neighboring octets.

This additional information about unit cell octet-truss structures provides necessary information to create a unit cell library to construct macro level component geometries.2.2 Octet-truss structure analysisThe octet-truss structure (displayed in Figure 1b) was chosen as the building block for the unit cells because of a previous analysis of the structure [2] and its good mechanical properties under most types of loading conditions. The initial analysis of the octet-truss structure was performed by Deshpande, Fleck, and Ashby [2] where their analysis rigorously studied the octahedral portion (Figure 2a) of the octet-truss structure for stiffness, compliance, and buckling. They state that the octahedral portion of the octet-truss structure will contribute the majority of the stiffness of the structure under compressive loading.

They state that the octahedral portion of the octet-truss structure will contribute the majority of the stiffness of the structure under compressive loading. Therefore their analysis is restricted to the octahedral section of the octet-truss structure and does not include the tetrahedral section of the structure (black section displayed in Figure 2b).Herein, a further analysis of the octet-truss structure is presented that includes the eight tetrahedrons of the octet structure for compliance. Both of these algebraic presentations consider each strut to be pinned at their respective nodes. An algebraic formulation for struts that are fully constrained at each node becomes too complex when all of the loading conditions (tensile, bending, shear, and torsion) are included and will not be presented.(a)(b)Figure 2.

(a) Octahedral section of the octet-truss structure (red inner section) and (b) octet- truss structure with highlighted a corner tetrahedron section (black section).The relative density of a unit cell will be used as a metric for characterization of the mechanical behavioral properties of the cell. The unit cell’s relative density is determined by the thickness of each octet-truss strut and number of octet-truss structures located within the cell defined by2a l(cid:167) (cid:168) (cid:169) where a is the strut diameter and l is the length of each strut. Equation 1 represents the material volume of the inscribed unit cell containing an octet-truss structure, specifically, only half of the volume of each strut on each of the six faces is included because the remaining volume lies in the volume for adjacent cells.

Each unit cell is analyzed for its stiffness (compliance) with respect to(cid:183) (cid:184) (cid:185)62(cid:83)(cid:85) (cid:32)423(1)the relative density. Unit cells containing n x n x n arrays of octet-truss structures are also be analyzed (for n = 1,2,3).Previous investigations of the octet-truss structure by Deshpande, Fleck, and Ashby [2] assumed that each node of the structure was pinned and free to rotate about the node, thus assuming only tensile conditions are present within the struts (struts are either in tension or compression). Additionally, their analysis primarily involved investigation of failure of the octet- truss structures due to elastic buckling or plastic collapse.

Additionally, their analysis primarily involved investigation of failure of the octet- truss structures due to elastic buckling or plastic collapse. Thus, bending moments, shearing forces, and torsion effects were not included within their analysis.2.2.1 Octahedral structure analysisDeshpande, Fleck, and Ashby [2] have determined that the octahedral section of the octet- truss structure (displayed in Figure 2a) determines the stiffness for the entire truss structure. For their analysis, they assume that all of the joints are pinned and only axial forces are present within the struts of the octahedral section.

(Note that different than printed in [2], we believe that there is a typographical error in the publication.) The compliance matrix [C] stated in Equation 3 is for a pinned node octahedral structure (Figure 2a) and not the entire octet-truss structure (Figure 2b). Deshpande’s analysis uses the compliance matrix [C] to determine the stiffness of the octet-truss and then makes a comparison to FEA results.

Deshpande’s analysis uses the compliance matrix [C] to determine the stiffness of the octet-truss and then makes a comparison to FEA results. The eight tetrahedral structures that are required in conjunction with the octahedral structure to create the octet-truss structure are not included in these calculations and are not represented in Equations 2-4.1 E16 E (cid:85)(cid:167) (cid:168) (cid:169)(cid:183) (cid:184) (cid:185)s(cid:32)(cid:32)3ss3424(2)(3)(4a)(4b)(4c)2.2.2 Pinned octet truss structure analysisThe authors concur with Deshpande that the majority of the compressive or tensile loads are absorbed within the octahedral section of the of the octet-truss structure, but the tetrahedral elements at the corners also contribute to the compliance or stiffness of the complete octet structure.

This is especially true when analyzing arrays of octet-truss structures, which will be illustrated in the following sections.The direct stiffness method can be used to determine the compliance matrix and reproduce Equations 2-4 for a pinned node octet-truss structure [3].

The coefficient fors and2s decreased from 9 to2649.8(octahedral to octet structure) and the(cid:124)2 decreased from 3 to , both decreasing by ~5%). The decrease in the compliance coefficients implies that the octet structure is less compliant (or more stiff) than the octahedral structure (along the principle load axis) and in transverse directions under loading (see Figure s did not change between the octahedral and octet analysis because the 6a). The shearing term twisting effect of the octet truss structure was not constrained to remain planar under the applied shearing load (i.e., the corner nodes were allowed to move out of plane due to the twisting nature of the shearing force).

Since this was only an analysis of a single octet, the shearing effect will become more significant when the octet structures are combined to create arrays of octets (i.e., unit cells).12283.2(cid:124)32.3 Unit truss Finite Element Analysis (FEA) approachUnit cells comprised of octet truss structures are currently being analyzed using a unit-truss finite element analysis program in MATLAB that has been developed at Georgia Tech [1]. The unit-truss finite element program is capable of non-linear analysis, but for the current stage in this project, only linear elastic behavior is considered. The FEA program includes axial, bending, shearing, and torsion effects that are present in loading of the octet truss structure where previous analyses only included axial effects [2].

The FEA program includes axial, bending, shearing, and torsion effects that are present in loading of the octet truss structure where previous analyses only included axial effects [2]. A discussion of the effective mechanical behavior of the unit cells using the unit truss FEA program will be discussed.425(6a)(6b)(6c)sIn the unit truss approach, a unit truss is used as a new unit cell for mechanics analysis of cellular structures, including lightweight structures, and compliant mechanisms. A unit truss consists of a central node and a set of half-struts that are connected to the node. Every two neighboring unit trusses share a common strut. An example of unit trusses is shown in Figure 3.Unit Truss 1Unit Truss 2Unit Truss 3Figure 3. Series of three unit truss structures that are connected between each node.Unit trusses can be parameterized, analyzed, patterned, and manufactured to support the desired design.

Series of three unit truss structures that are connected between each node.Unit trusses can be parameterized, analyzed, patterned, and manufactured to support the desired design. In a unit truss, the strain and stress around the nodes (displayed in Figure 4), are usually complicated due to considerable inter-strut interactions and large bending moments [1]. The unit truss is leveraged from the ground truss approach and homogenization method [5, 6]. The constitutive equations of 2-D and 3-D unit trusses are shown in Equations 7-9. The linear elasticity of a unit truss is represented by eK , while U represent the nodal displacements (cid:5) (cid:5) and forces. Unit trusses can have any number of incident struts and they are special finite elements for analyzing large cellular structures.and F (cid:5)u (2) 1u (1) 1u (1) 3u(2) 2u (2) 3u(1) 2u (0) 1u(0) 2u(3) 2u (3) 1u (5) 3u (0) 3u (3) 3u (5) 1u(5) 2u (4) 3u (4) 1u(4) 2Figure 4.

Using the unit truss approach, combined with the tangent stiffness method, truss structures can be analyzed under axial forces, bending, torsion, nonlinearity, and buckling [7].2.4 Unit cells comprised of octet truss structuresThe unit cell analysis will be applied to three different types of unit cells that are comprised of octet-truss structures, displayed in Figure 5. The analysis is based on the approach used by Deshpande by relating the relative stiffness (E/Es) to the relative density ( (cid:85) in Equation 1) for a fixed sized unit cell. The relative density for a unit cell geometry is altered by varying the diameter of the struts of the octet-truss structure. Unit cells comprised of a single octet, a 2x2x2 array, and a 3x3x3 array are analyzed within this work, but this analysis can easily be expanded to arbitrary unit cells and octet arrays.(b) Figure 5.

CAD images of unit cells comprised of octet-truss structure, (a) single octet, (b) 2x2x2 array of octets, and (c) 3x3x3 array of octets.(a)(c)2.5 Unit truss FEA analysis of unit cellsA direct comparison of the relative stiffness of unit cells that are composed a single octet and an array of octet-truss structures is presented herein. To maintain uniformity between analyses of the different mesostructures (unit cells constructed from an array of octets), the dimension of the unit cell is set to a fixed value and the applied compressive load is appropriately distributed to each node creating the same uniform applied stress. A comparison between the unit-truss FEA approach (described in Section 2.3) of a mesostructure unit cell and Deshpande’s analysis of the octet-truss structure is detailed below.

A comparison between the unit-truss FEA approach (described in Section 2.3) of a mesostructure unit cell and Deshpande’s analysis of the octet-truss structure is detailed below. Only linear elastic deformations have been considered in this work, but the unit-truss FEA program is capable of determining non-linear deformation and buckling effects, which will be analyzed in future work.Figure 6 displays the relative stiffness plotted with respect to the relative density of the mesostructure unit cell. An increase in the relative stiffness is observed as the relative density increases for the unit cells. To increase the relative density of the unit cell is to increase the strut diameter, and a larger diameter strut will possess a greater stiffness value that translates to the overall relative stiffness of the structure.

Log-log plot of the relative stiffness (E/Es) verses relative density for (a) the pinned node octahedral (Eq. 4) and octet structure (Eq. 6), and (b) the pinned node octet structure and the unit-truss FEA analysis of unit cells (analysis includes tensile, shear, bending, and torsion).The unit-truss FEA results of unit cells comprised of a single, 2x2x2, and 3x3x3 array of octet truss structures is displayed in Figure 6b. The general trend of an increase in relative stiffness (E/Es) is observed, but the unit-truss FEA results indicate that the rate of change in relative stiffness with respect to relative density is less than predicted by the algebraically determined relative stiffness (Equations 4 and 6). A comparison between Eq. 6 (valid for only a single octet) and the unit-truss FEA results for a single octet indicate that Eq.

A comparison between Eq. 6 (valid for only a single octet) and the unit-truss FEA results for a single octet indicate that Eq. 6 predicts higher values of relative stiffness for03.0(cid:33)(cid:85), as displayed in Figure 6b.The unit-truss FEA results also indicate that as the number of octet truss structures increase within a fixed unit cell geometry, the relative stiffness also increases. Specifically, a unit cell with a fixed relative density has a greater relative stiffness with 27 (3x3x3) octet-truss structures (Figure 5c) when compared to the unit cell containing eight (2x2x2) octet-truss structures (Figure 5b). Therefore, unit cells comprised of octet-truss structures with a fixed relative density becomes more stiff (less compliant) when the number of octet-truss structures increases. In order to increase the number octets within the unit cell and still maintain the same relative density, the strut diameter and length must both be decreased.

In order to increase the number octets within the unit cell and still maintain the same relative density, the strut diameter and length must both be decreased. Indicating that the relative stiffness is more sensitive to the strut length than the strut diameter (by Equation 1). Unfortunately, this result was not observed experimentally, which is presented in Section 3.3 Experimental compression tests of unit cells comprised of octet-truss structuresUnit cells comprised of octet-truss structures were fabricated and subjected to a mechanical compression test to determine their material properties. The unit cells were comprised of a single, 2x2x2, and 3x3x3 array of octet-truss structures, as displayed in Figure 5. These unit cell test specimens were created to observe the effect on the unit cell’s mechanical properties with respect to the number of composing octet-truss structures.

These unit cell test specimens were created to observe the effect on the unit cell’s mechanical properties with respect to the number of composing octet-truss structures. The compression test procedure for the unit cells followed the standard compression test method for determining compressive properties of rigid plastics, ASTM D 695-02a.42813.1Specimen design and fabricationA relative density value of 0.35 was chosen to compare the effect of changing the number of octet-truss structures in the unit cell. The test specimen unit cell was designed to be a 50 mm cubic cell, containing n3 octet-truss structures (n = 1, 2, or 3). In order to create geometric conditions as close to the octahedral analysis conducted by Deshpande [2], the octet-truss structures were arranged such that there are either 1, 8, or 27 octahedral structures in the unit cell (formed by the octet-truss structures). The geometry of the unit cells are displayed in Figure 5.

Material properties of RenShape™ SL 7510 resin.Material PropertyValueTensile mod. (Es) Tensile Str. ((cid:86)Y) Elongation at break Density2.634 [GPa] 57 [MPa] 10.1 % 1.18 [g/cm3]The experiment specimens were manufactured on a 3D Systems SLA3500 using RenShape™ SL 7510 resin. The “exact” build style was used for specimen fabrication with a layer thickness of 0.1016 mm (0.004 in) and a beam width of 150 (cid:80)m. The samples were ultrasonically cleaned in isopropyl alcohol for 10 minutes and post cured in a UV oven for 60 minutes. The samples were isolated from UV light and sealed in a plastic bag to minimize any environmental effects on the specimens before testing. The samples were tested approximately 40 days after fabrication.3.2 Empirical compression measurements and resultsThe unit cell specimens with relative density of 0.35 exhibited both brittle and ductile failure responses.

The unit cells containing a single octet and the 2x2x2 octet array both resulted in catastrophic failure with little or no audible warning before failure (complete destruction). The stress-strain curve (displayed in Figure 7) clearly shows a brittle failure response for the single octet and 2x2x2 array unit cell. The 3x3x3 octet array unit cell exhibited a more ductile failure accompanied by audible cracking before yielding. Some of the 3x3x3 octet arrays did not catastrophically fail and these specimens remained relatively intact after testing. The initial behavior of the 2x2x2 and 3x3x3 unit cells were very similar possessing almost the same stiffness (0.060 and 0.065 GPa, respectively), but the 2x2x2 had a brittle failure response while the 3x3x3 sample exhibited a more ductile failure response.

The average experimental results for the unit cell octet arrays is provided in Table 3.The amount of strain produced in the unit cells increased as the number of octet-truss structures increased, specifically from (cid:32)x3(cid:72) 0.0889 for the 3x3x3 unit cell. Ironically, the same trend can not be stated for the maximum allowed stress and unit cell stiffness. The maximum applied stress and the stiffness for the 2x2x2 octet array unit cell are less than the values for both the single and 3x3x3 octet array unit cells. The stiffness of the(cid:32)x1(cid:72) 0.0723 for the single to4293x3x3 octet array unit cell (Eemp=0.065 GPa) was significantly less stiff than the single octet unit cell (Eemp=0.107 GPa) and required approximately half of the applied load (Fmax) to achieve 18% more strain than the single octet unit cell.Table 3.

Stress strain plot of the compression tests for the unit cells comprised of octet-truss structures (all possessing a relative density of 0.35) and the ASTM specimens.The single octet unit cell possessed the greatest relative stiffness (E/Es) of 0.0405 (or approximately 4% of Es) while the 3x3x3 possessed approximately half the stiffness of the single octet with a relative stiffness of 0.0247 (or 2.47% of Es). All of the unit cell specimens experienced total displacements ranging 4.00-4.62 mm resulting from maximum loads ranging from 8.7-15.4 kN.3.3 Unit truss FEA computations for unit cells comprised of octet-truss structuresA FEA simulation of the compression experiments conducted in Section 3.2 using the unit truss FEA program is presented in this section for the octet-truss structures (unit cells).

The FEA analysis has determined that the 2x2x2 octet unit cell is the most compliant (least stiff) of the octet-truss unit cells and allowing the largest amount of strain to be produced.3.4 Comparison of analytical, experimental results, and ,unit truss FEA computationsA comparison of the analysis approaches of the octet-truss unit cells presented in Sections 2.2, 3.2, and 3.3 is displayed in Table 5. The relative stiffness equations (Eq. 4 and 6) are for a single truss structure (octahedral and octet, resp.) and are also used for unit cells comprised of arrays of octahedral/octet truss structures (Deshpande [2] also employed this type of scaling).Table 5. Relative stiffness comparison between the analytical octahedral, octet, and unit-truss FEA compliance to the empirical relative stiffness results for a unit cell with35.0(cid:32)(cid:85).Octet Array SizeSingle 2x2x2 3x3x3(Table 3)Eemp Es0.0405 0.0228 0.0247(Eq. 4) Eoctahed Es 0.0389 0.0389 0.0389(Eq.

The unit-truss FEA was very consistent in is predictive capability of the relative stiffness with relative errors ranging tightly about 10% for all of the unit cells analyzed. No significant accuracy was lost in changing the number of truss structures comprising the unit cells.4 ConclusionsAn analysis of unit cells comprised of different sizes of octet truss structures has been performed. An analytical formulation for the octet-truss structure has been presented that can more accurately predict the relative stiffness for a single octet truss structure, which has been confirmed experimentally. A unit-truss FEA program has been employed to simulate compression testing of unit cells, exhibiting an accuracy of 10% relative error to experimentation431for the relative stiffness of all of the unit cells analyzed.

The unit truss FEA program shows promise as an analysis tool for unit cells comprised of arbitrary truss structures.5 Future research tasksThe analysis of the octet-truss structure is the beginning of the next few steps for the mesostructure design synthesis. A library of unit cells will be created after determining the behavior of various n x n x n arrays of octets under different loading conditions (non-uniform loads, shearing loads, and torsion). These tasks may now progress with the knowledge that the unit-truss FEA program is capable of simulating empirical results.The library of the mesostructure unit cells can then be used as building blocks for component geometries. An optimization program can be applied to design the component geometry for lightweight by using less dense unit cells with either greater or less stiffness where applicable due to defined loading on the component.

For each loading condition, an optimization function will be used to reduce internal energy, increase strength, and reduce weight based on the application of the structure. The optimization function will accomplish this through the removal of unnecessary elements, increasing the cross-sectional area of struts with high orders of stress, and reducing the cross-sectional area of those with low orders of stress. The goal in this is to create a component comprised of unit cell truss structures that have been designed according to their required strength, flexibility, toughness, etc.6 AcknowledgementsFinancial support for this work has been funded by the National Science Foundation (NSF) grant DMI-0522382. The authors would like to thank the following individuals at Georgia Tech for their individual contributions to this research: Robert Cooper for helping with compression testing, and Jeff Lloyd and Amanda O’Rourke for CAD and image generation.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperANALYSIS OF POWDER REMOVAL METHODS FOR EBM MANUFACTUREDTI-6AL-4V PARTSAmit J. Lopes1,3, Luis C. Ramos3, David Saenz1,2, Philip Morton1,2, Cesar A. Terrazas1,2, Ahsan Choudhuri2, and Ryan Wicker1,2.1 W.M. Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, TX 79968, USA 2 Department of Mechanical Engineering, The University of Texas at El Paso, El Paso, TX 79968, USA 3 Department of Industrial, Manufacturing and Systems Engineering, The University of Texas at El Paso, El Paso, TX, 79968, USAAbstractAdditive Manufacturing (AM) allows the creation of complex geometries that are not achievable through subtractive manufacturing. Regardless of the advantages that 3D Printing offers, technology limitations often constraints the desired geometry.

Regardless of the advantages that 3D Printing offers, technology limitations often constraints the desired geometry. When fabricating Ti-6Al- 4V parts in Electron Beam Powder Bed Fusion (EBPBF), the electron beam is used to preheat the powder bed to maintain the desired temperature gradient during the build. One disadvantage of EBPBF during the fabrication process is the trapped powder within internal channels gets partially sintered and require extra processing. This research analyzes several powder removal methods and compares their effectiveness. This work utilizes two types of samples, both made of Ti-6Al- 4V in EBPBF; with geometries that resemble typical features when designing a component. The target weight of each cylinder is calculated based on dimensions and effective density of the sample.

The target weight of each cylinder is calculated based on dimensions and effective density of the sample. The results summarizing the effectiveness of each method are presented.IntroductionWhen fabricating parts using Electron Beam Powder Bed Fusion (EBPBF), the powder bed is preheated using the electron beam; Gong et al. indicates that this serves two purposes: holding the metal in place during the subsequent melting scan and reducing the thermal gradient in the build part [1]. The decrease in temperature difference between layers helps to avoid residual stresses within the printed part, making it one of the advantages of this technology. Nevertheless, one of the main disadvantages of EBPBF is the sintered powder that gets trapped inside of the build part during the fabrication process.

Nevertheless, one of the main disadvantages of EBPBF is the sintered powder that gets trapped inside of the build part during the fabrication process. The trapped sintered powder inside of channel-like features is of primary interest as 3D printed components often have lines and channels that require a clear passage for their end design.Studies and experiments have been conducted to remove powder from parts made using Selective Laser Sintering (SLS), Tenney et al. (2018) proposed the use of traveling waves as a de- powdering process for additively manufactured parts [2]. Thompson et al. (2015) used venting holes in the prototype design to remove the packed powder via pumped fluid; these vent holes were integrated to de-powder each layer separately [3]. The sintered powder is not an issue on laser-based systems since most of the processing is done at low temperatures [4].

The sintered powder is not an issue on laser-based systems since most of the processing is done at low temperatures [4]. In the case of EBPBF, few studies have been conducted to research methods to remove the sintered powder.2064Hasib et al. (2015) presented a method which consisted of exposing an EBPBF printed Ti-6Al-4V lattice to a chemical etcher with results that are not completely successful and conclusive [4].The purpose of this research is to perform preliminary analysis of potential methods that can remove the sintered powder from channels and hollow features from parts made of Ti-6Al- 4V, manufactured with EBPBF. Figure 1 shows the two types of cylinders used for this study; a cylinder with straight channels and a cylinder curved channel. The selected geometries resemble typical channel design configurations used in applications such as cooling lines.

The selected geometries resemble typical channel design configurations used in applications such as cooling lines. The methods to evaluate are Powder Recovery System (PRS), Ultrasonic, Ultrasonic with Mechanical Impact, Vapor Blast, Chemical Etching, and Liquid Nitrogen with Ultrasonic. The samples were printed on an ARCAM A2X machine with ARCAM’s default standard parameters for Ti-6Al-4V alloy.0-.¢ /_ Q ., o . ·--~- --.,{f,""'<t;~ \;,(a)(b)(c)Figure 1 – (a) CAD cylinder design with straight channel features, (b) CAD cylinder design with curved channel features, and (c) Additive Manufactured Cylinder.MethodologyTo evaluate the powder removal methods, fourteen parts; seven curved- channel cylinders and seven straight-channel cylinders; were manufactured to evaluate each proposed methodology. All parts underwent the Powder Recovery System (PRS) methodology before any of the proposed research methods were applied.

All parts underwent the Powder Recovery System (PRS) methodology before any of the proposed research methods were applied. Each described research method was used on one curved-channel cylinder and one straight-channel cylinder. All of the samples were weighed before and after all powder removal methods were applied. To facilitate this comparison, the effective density of EBPBF part was calculated, as described in the next section.Effective DensityThe density of uncured Ti-6Al-4V powder is estimated to be 2.54 g/cm3 while the density of Ti-6Al-4V ranges from 4.429 g/cm3 – 4.512 g/cm3. To determine a value that reflects the as- built density of a part made of Ti-6Al-4V using the EBPBF process, the mass of three printed solid cylinders was measured and averaged. Each part is a cylinder 25 mm in diameter and 25 mm height and is illustrated in Figure 2.

Each part is a cylinder 25 mm in diameter and 25 mm height and is illustrated in Figure 2. After the build, each part was individually cleaned with the PRS to remove all sintered powder from its walls, measured to determine its volume, and weighed to calculate its corresponding density. The average of the three calculated values resulted in an as-built effective density of 4.2691 g/cm3.

The average of the three calculated values resulted in an as-built effective density of 4.2691 g/cm3. This calculated density was used in the CAD mathematical model to have the CAD software calculated mass reflect a result, more accurate to a real printed component.2065Density = 4.27 grams per cubic centimeterMass = 52.39 grams(a)(b)(c)Volume = 12.27 cubic centimetersFigure 2 – (a) Solid cylinder build to calculate the density of Ti-6Al-4V after EPBF, (b) CAD model of the cylinder, and (c) the mass properties results from SolidWorks 2018.Powder Recovery System (PRS)The first methodology identified for the powder removal process, to evaluate its effectiveness was the Powder Recovery System (PRS), which is an auxiliary machine that is supplied with the Arcam EBM machine. The PRS operates under a pressure of 6 mBar to pressurize air and Ti-6Al-4V powder, which is required to breakdown sintered powder and facilitate its removal.

The energy is transferred with the use of an attachment which transfers the energy onto a part using a frequency that impacts the part. The apparatus operates at the same frequency of 20 kHz at all times while the amplitude can be varied. For this methodology, the amplitude was kept constant 40% power. Using a process similar to that of the PRS, each sample parts is affected by the Ultrasonic process for 2-minutes and then the post process weight of the part is recorded. This process is repeated for several intervals until the parts are either cleared of the trapped powder or until no significant weight difference is observed from the previous interval.Ultrasonic and Mechanical Impact Powder RemovalThis method adds a mechanical impact process using a hammer in addition to Ultrasonic method previously described. The hammer or mallet used has a head made of rubber on one side and plastic on the other side to prevent the parts from getting damaged due to repeated impacts.

The hammer or mallet used has a head made of rubber on one side and plastic on the other side to prevent the parts from getting damaged due to repeated impacts. The Ultrasonic and mechanical impact approach required to mechanically impact the part for one minute (to break off any loose powder), followed by the Ultrasonic for 2 minutes to further remove trapped powder inside of each cavity, and finally mechanically impact the part for thirty seconds.2066The 2-minute ultrasonic approach was followed to maintain consistency with the rest of the testing procedures.Figure 3 – (Left) Ultrasound being applied onto the part. (Right) Part being hammered.Vapor Blast Powder RemovalThe Vapor Blast (VB) testing is a powder removal method that uses a machine that cleans the surface of metal parts using a slurry of water and sand. Inside the machine chamber, the pressurized slurry is sprayed through a nozzle and directed to the cylinder channels.

Inside the machine chamber, the pressurized slurry is sprayed through a nozzle and directed to the cylinder channels. This testing process is different compared to the previous methods since it involves water. Each part of the sample pair was cleaned using the VB process for five minutes intervals, after which the parts were dried inside an oven for about 40 minutes at 120°C. After the drying process is complete, the parts were weighed, and the values were recorded. The testing process was repeated for the required iterations until the current iteration shows no significant weight difference.Chemical Etching Powder RemovalPowder removal using chemicals used for etching; Kroll’s reagent and Keller’s reagent, was another method evaluated as part of this research. Chemical etching is a technique used to remove material using chemicals where the corrosive chemical, known as the etchant, reacts to the target material, dissolving the material at a time-dependent rate.

Etchants, in general, are used for the dissolution of the surface of a metal part and are mostly used for metallography to analyze the microstructure. Kroll’s reagent is an etchant used for Ti-6Al-4V, this solution is 92 mL of deionized water, 6 mL of nitric acid, and 2 mL of hydrofluoric acid. Keller’s reagent is an etchant used for aluminum, Ti-6Al-4V has 6% aluminum, making this etchant another viable method to analyze. The solution is 90 mL of deionized water, 5 mL of nitric acid, 3 mL of hydrofluoric acid, and 2 mL of hydrochloric acid. The testing process for this method requires exposing a part to the etchant. The testing method is the same for both etchants and requires placing droplets of the solution into the holes of the part, waiting for one minute to allow the solution to affect the part, and observe if any additional powder is removed. Then, the part is rinsed and dried using an oven before weighing the part.

Then, the part is rinsed and dried using an oven before weighing the part. The testing process was repeated for the required iterations until the current iteration shows no significant weight difference.2067Liquid Nitrogen and Ultrasonic Powder RemovalThe final powder removal method analyzed used a combination of ultrasound and liquidnitrogen. The first step involved submerging the sample parts (with curved and straight holes) into liquid nitrogen for twenty minutes, followed by an ultrasonic application for two minutes. The parts are weighed, and their values recorded, and the testing process repeated until the current iteration shows no significant weight difference from the previous one. Figure 4 shows the parts submerged in liquid nitrogen.Figure 4 – (Left) Parts before they are submerged inside a pot with liquid nitrogen. (Right) Parts submerged inside a pot with liquid nitrogen.ResultsThe results of each methodology are shown in a table for comparison.

(Right) Parts submerged inside a pot with liquid nitrogen.ResultsThe results of each methodology are shown in a table for comparison. The table presents the type of cylinder, its weight before the PRS, its weight after the PRS, and the weight difference after being exposed to the assigned methodology. An estimate evaluation of the theoretical weight of the parts and the target weight was conducted to evaluate the effectiveness of each method.Weight ComparisonThe first analysis performed included all of the information on the PRS cleaning. Thecleaning of all the samples with the PRS allowed gathering more sample information on this methodology.

Thecleaning of all the samples with the PRS allowed gathering more sample information on this methodology. Table 1 below shows the average weight before and after for 14 samples.Table 1 – Table Summary of Powder Recovery System (PRS) MethodProcedure PRS (7 Set average)Weight Change (grams) After (grams) Straight Curved Straight Curved Straight Curved 2.9236 112.9015 112.7740 109.9842 109.8504Before (grams)2.9173After all the sample cylinders were clean with the PRS, a pair of cylinders, consisting of one straight channel cylinder and one curved channel cylinder, were exposed to the assigned methodology.

There are non-destructive evaluation methods that can assist in the effectiveness evaluation, for this preliminary work, it was decided to use mathematical methods to calculate a rough estimate on the mass and volume of each cylinder before any method was applied.Target WeightDuring the part formation in the Electron Beam Powder Bed Fusion, the powder goes through phase changes, from powder to liquid and then solid. One main advantage of building in an Electron Beam machine is the preheat cycle that reduces the temperature difference of the powder between layers. This added step in the process is an excellent aid in reducing the residual stresses, but when the part is finally cooled down and removed from the machine, the final part does experience a degree of shrinkage.

To understand how much powder needs to be removed from a newly printed part, the CAD mathematical model was taken as a reference point of start and adjusted its density to the effective density initially calculated to provide an ideal weight. From the software calculations, the initial ideal weight was found to be 117.4871 grams for the cylinder with straight channel cavities, and the ideal weight for the cylinder with curved channel cavities was found to be 116.9109 grams.The ideal weights represent a CAD model with no change in dimensions caused by shrinkage. The actual printed part experiences a degree of shrinkage, and physical measurements were done to verify the change in volume. The exterior dimensions were easily measured, whereas the 2mm & 1mm holes represented a challenge to measure with simple tools. A trend line analysis was used to evaluate the shrinkage for the 2mm & 1mm holes, using the holes actual dimensions and the actual outer diameter of the cylinder.

With the actual dimensions, now it is possible to calculate the volume of printed cylinders. Table 4 below shows the ideal theoretical weight and the calculated target weight.Table 4 – Actual Post-Build Weight Based On Shrinkage CalculationsWeight (grams)Part massMathematical Model Curved 116.9100Straight 117.4900Actual build partsStraight 107.4900Curved 107.1200EffectivenessThe calculated target weights were for reference use only and there are better methods to obtain the actual volume and mass of the cylinders, such as a CT scan. However, for this preliminary work, estimates were used to get a reference mass to use as a target weight. For the cylinder with straight channel cavities the target weight was estimated to be 107.49 grams, and the target weight for the cylinder with curved channel cavities was estimated to be 107.12 grams.

Each methodology was evaluated for its effectiveness with these target weights; Table 5 below shows the efficiency of each methodology for removing sintered powder after the PRS process.Result DiscussionThe PRS was the most effective method to remove sintered powder when the part was taken out of the machine. The PRS can remove only a portion of sintered powder, after a certain number of iterations, the change in weight was no different from the previous iteration. These proposed methodologies aim to remove the remaining sintered powder trapped in the features inside the printed geometry. From the methodologies tested, we see that the most efficient methods are: Ultrasonic, Ultrasonic & Mechanical Impact, and Liquid Nitrogen & Ultrasonic. Further research and testing of the most effective methodologies are to be conducted to improve the powder removal and, simultaneously, have more accurate information about the effectiveness of the methods.

The trapped sintered powder inside of a printed component is one of the many drawbacks that this Additive Manufacturing technology faces and little work have been done to understand how to remove the sintered powder. It is true that during the building process the sintered powder plays a vital role as support for the final printed component but imposes design limitations that do not allow to have the full advantage of this Additive Manufacturing technology. From the results obtained, the PRS is the most effective method to remove the sintered powder; however, after several iterations, the effectiveness comes to a stall. Ultrasonic, and Liquid Nitrogen & Ultrasonic are methods that can remove more sintered powder from the channels after the PRS. Ultrasonic waves can travel through the part, allowing it to reach the channels, while liquid nitrogen can easily access the channels directly and adapt to the channel form.

Ultrasonic waves can travel through the part, allowing it to reach the channels, while liquid nitrogen can easily access the channels directly and adapt to the channel form. This preliminary work shows that there are methods that can improve the removal of the sintered powder.Future WorkFuture work will involve the actual testing of the most promising methodologies obtained in a physical complex fuel injector. Additionally, it is intended to use a CT scan as a nondestructive method to evaluate the powder removal, instead of using a weight comparison method. The methodologies discussed in this paper could be used to analyze different channel geometries, e.g., square or triangular channels. Lefky et al. have performed research using chemical corrosion attack to remove support structures from metal printed components [6][7].

Lefky et al. have performed research using chemical corrosion attack to remove support structures from metal printed components [6][7]. The chemical corrosion attack or chemical etching is a method that will be further investigated since the etching rate is dependent on several factors, and the preliminary experimentation to determine this rate was not performed for this work. There are several chemical etchants more adequate for Ti-6Al-4V, and to determine the etch rate before any exposure to the part is vital for the methodology to work correctly. Additionally, the determination of how much surface is required to be removed by the chemical etchant to clear channels of the sintered powder is part of the future work to be done.A second experiment can be the evaluation of these proposed methodologies with parts that have were not initially exposed to the PRS process.

All of this preliminary research is the first step to address the need for further investigation in the topic of removing sintered powder from internal cavities on EBM manufactured parts.Special ThanksSpecial thanks to the W.M. Keck Center for 3D Innovation at The University of Texas atEl Paso for the use of their equipment and the personnel assigned for testing.2072References[1] Gong, X., Anderson, T., & Chou, K. (2014). Review on powder-based electron beam additive manufacturing technology. Manufacturing Review, 1, 2. Retrieved April 16, 2019.[2] Tenney, C. M., Malladi, V. V., Musgrave, P. F., Williams, C. B., & Tarazaga, P. A. (2018). Traveling Waves As a De-Powdering Process for Additively Manufactured Parts. Volume 1: Development and Characterization of Multifunctional Materials; Modeling, Simulation, and Control of Adaptive Systems; Integrated System Design and Implementation. Retrieved January 4, 2019.[3] Thompson, S. M., Aspin, Z.

alysis of Sealing Methods for FDM-fabricated PartsJorge Mireles1, Arturo Adame1, David Espalin1, Francisco Medina1,Rob Winker2, Terry Hoppe2,Bob Zinniel2, and Ryan Wicker11. W.M. Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, TX 2. Stratasys, Inc., Eden Prairie, MNAbstractAs a result of the layer-by-layer deposition characteristics of Additive Manufacturing (AM) processes, fabricated parts exhibit limiting qualities and have yet to achieve the requirements for end-use applications. Specifically, the use of AM-fabricated parts in fluid pressure applications is limited due to part porosity as well as non-optimized building variables (e.g., build orientation and material properties). In an effort to extend the use of AM in more applications involving fluid pressure, parts manufactured with Fused Deposition Modeling (FDM) were sealed with a variety of sealants and tested under applied pressure.

Eleven sealants with diverse chemical properties were applied to multiple geometries of FDM-fabricated pressure caps through brushing or vacuum infiltration. The caps were installed on pressure vessels and subsequently tested while safety precautions were taken to avoid catastrophic failure (i.e., exploding) caused by pressure differentials. Results of the testing provides a sealing method using BJB TC-1614 that enables FDM-fabricated parts to withstand pressures up to ~276 kPa (40psi) through brushing and ~138 kPa (20 psi) through vacuum infiltration. Other noteworthy sealants (Minwax Sanding Sealer, Minwax Polyurethane Oil Based, PRO Finisher Water-Base Polyurethane) that are readily available to consumers and easy to apply (i.e. no mixing ratios to follow, long working times) also had notable results by withstanding pressures up to ~207 kPa (30 psi).

no mixing ratios to follow, long working times) also had notable results by withstanding pressures up to ~207 kPa (30 psi). In addition, an analysis on dimensional changes was performed to determine the absolute difference between as-built and surface-treated parts. Parts that were infiltrated with BJB TC-1614 showed less dimensional changes (average absolute change of 0.104 mm) than parts that were brushed (average absolute change of 0.231 mm) however one- part sealants had smaller dimensional changes (maximum absolute change for one-part sealants of 0.065 mm for infiltration and 0.171 for brushing) with noteworthy results in pressure testing. Benefits of filling voids within FDM-manufactured parts enables end-use applications such as hermetic housings for biomedical devices and pipes/covers for thermodynamic systems such as heat exchangers.1 IntroductionFused Deposition Modeling (FDM), developed by Stratasys, Inc.

(Eden Prairie, MN), is a solid freeform fabrication or Additive Manufacturing (AM) technology that can potentially build end-use parts through the deposition of materials layer-by-layer without the need for part- specific tooling or molds, which reduces the cost and lead-time for low-volume production. Production-grade thermoplastic materials such as acrylonitrile butadiene styrene (ABS), methyl methacrylate acrylonitrile-butadine styrene (ABSi), polyphenylsulfone (PPSF), polycarbonate (PC), and other blends of ABS (ABS-M30, ABS-M30i, ABS Plus, PC-ABS) are available for use with FDM technology [Chua et al., 2003]. Parts produced by an FDM 400mc (Stratasys, Inc., Eden Prairie, MN) exhibit an accuracy of ±0.127 mm (0.005 inch) which is useful for building high resolution parts [Chua et al., 2003].

Parts produced by an FDM 400mc (Stratasys, Inc., Eden Prairie, MN) exhibit an accuracy of ±0.127 mm (0.005 inch) which is useful for building high resolution parts [Chua et al., 2003]. Although FDM technology has been used in a variety of end-use part applications [Wohlers, 2005], parts produced by FDM cannot be used influid pressure applications because they are generally porous [Agarwala et al., 1996; Bellini et al., 2004]. In addition, the FDM building process produces moderately strong bonds between layers; however, voids at these locations poses a safety hazard.

In addition, the FDM building process produces moderately strong bonds between layers; however, voids at these locations poses a safety hazard. That is, parts exposed to elevated fluid pressures can give rise to a pressure differential and produce catastrophic failures (e.g., exploding).Sealing of FDM parts is useful in applications such as low pressure heat exchangers (100- 150 kPa) for air dehumidification and waster recovery [Kuppan, 2000], hermetic housings for biomedical devices such as difibrillators or pacemaker [Nagl and Lechleitner, 2005] as well as some advanced applications for low pressure tanks that are resistant to impact and are able to withstand corrosive environments [Koppert and Beukers, 2000; Rashilla, 1993].

Implementing FDM-manufactured parts in such applications requires a post-process that eliminates voids and reinforces the bonds of FDM layers to produce an impermeable, end-use part with reduced safety hazards.Sealing options (brushing, solvent dipping, infiltration, raster modifications, vapor smoothing, dipping, and painting) have been used to seal parts with apparent porosities useful in applications of high strength and where wear-resistant surfaces are needed [Bunshah, 1994, Lee et al., 2004]. Previous research has shown vapor smoothing to improve surface finish of FDM- manufactured parts, and although work needs to be conducted to investigate its efficacy for sealing, vapor smoothing does not appear to affect part dimensions (mean difference between after and before smoothing of 0.028 mm) [Espalin et al., 2009]. Another method used for sealing is vacuum infiltration wherein a porous part is exposed to a fluid in a vacuum allowing the fluid to fill the pores within the part.

Another method used for sealing is vacuum infiltration wherein a porous part is exposed to a fluid in a vacuum allowing the fluid to fill the pores within the part. Manufacturing facilities not equipped with vacuum chambers may implement a simple brushing method to directly apply sealants by using foam or bristled brushes.In this study, two sealant application methods (vacuum infiltration and brushing of sealants) and eleven sealants were explored to fill air gaps left by the FDM manufacturing process. Experiments were conducted to determine if the post-process enhanced the FDM- manufactured parts to further allow for fluid pressure applications while minimizing dimensional changes caused by the sealant and/or application method.

Protocols developed in this study can be used as guidelines at the manufacturing level and results can aid FDM users in properly selecting a sealant/infiltrant.2 MaterialsABS-M30 (Stratasys, Eden Prairie, MN), was used with an FDM 400mc system to produce test parts described in section 3. ABS-M30 is a popular option for consumers because of its increased strength compared to regular ABS (25-70% stronger) and its range of available colors (ivory, white, black, dark grey, red, blue) [Chua et al., 2003; Wohlers, 2011]. Various sealant materials were used and are described in Table 1.3 Methods3.1 Design of multi-feature test partPressure caps and vessels are designed to hold fluids at pressures different than atmospheric pressure. Design of pressure vessels and caps vary and common shapes include spheres and cylinders [Spence et al., 1994] which were included in the multi-feature test part for this study.

Design of pressure vessels and caps vary and common shapes include spheres and cylinders [Spence et al., 1994] which were included in the multi-feature test part for this study. Prominent features that were implemented included cylindrical pressure vessels andTable 1: Densities of tested sealants and typical applicationsSealantDensity (g/cm3)Typical ApplicationsDEFT Clear Brushing LacquerIPS Weld-On 3 CementMinwax Sanding Sealer0.901.330.85Finish for interior wood surfaces. Cementing acrylic and other plastics; used in display items, containers, house wares, etc. Seals interior wood surfaces.Minwax Oil Base Polyurethane0.86Protects interior wood surfaces.PRO Finisher Water-Base Polyurethane1.00Coating hardwood floors and other interior wood surfaces.Thompson's Multi-Surface WaterprooferBJB TC-1614 A/BHysol E-30CLStycast W19 + Catalyst 9West Marine Penetrating Epoxy0.811.141.101.090.93Protects and seals wood, brick and concrete.

Seals porous to semi-porous material surfaces like plaster, wood, graphite and SLS models. Bonding, small potting, laminating glass, ceramics, metals, and plastics. Impregnates wrapped coils; small device potting; surface coating. Seals dry-rotted wood: repair transoms, decking, stringers and molding.West System 105 Resin + 209 Hardener epoxy1.15Coating and bonding applications in extremely warm and/or humid conditions for fiberglass, composite materials, and a variety of metals.their end caps, 90 degree elbows, and other design options such as square pressure containers and semi-cylindrical features. The final design of the multi-feature test part is portrayed in Figure 1 and includes an example application for each feature.Variable feature thicknesses were used to evaluate the effect of sealant permeability on part performance during pressure testing.

A detailed view of the multi-feature test part and its corresponding thicknesses is portrayed in Figure 2.90 degree fittingsquare pressure containerspherical end capsemi cylindrical pressure vesselFigure 1: Multi-feature test part designThe base thickness is 0.33 cm.The post on the top of the large dome and the squares attached to the small domes are solid.The large dome thickness is 0.64 cm.The small domes range in thickness: 0.43 cm, 0.48 cm and 0.53 cm.The four square reference geometries and elbow features range in thickness: 0.20 cm, 0.25 cm, 0.30 cm, and 0.35 cm.Figure 2: Top view of multi-feature test part thicknesses3.2 Vacuum InfiltrationTest parts were submerged in the sealant and exposed to a vacuum environment (0.375 torr) as shown in Figure 3 via an MCP 4/04 vacuum casting system (MTT Technologies Ltd., Whitebridge Way, Whitebridge Park, England). All tests and application of sealants were performed at ambient room temperature.

All tests and application of sealants were performed at ambient room temperature. Vacuum infiltration is a process which consumes a large portion of infiltrant depending on part size and area of infiltration. To avoid material waste, the manufacturer-recommended sealant work times and mixing ratios were used. An infiltration study was performed on small test specimens (76.20 X 19.05 X 3.18 mm) and their cross-sections were observed through the use of a stereo microscope to verify the part was being infiltrated with each sealant. Based on this initial infiltration study, candidate sealants were identified for vacuum infiltration of the multi-feature test part in Figure 1.

Based on this initial infiltration study, candidate sealants were identified for vacuum infiltration of the multi-feature test part in Figure 1. These sealants included Minwax Sanding Sealer, Polyurethane Oil Based, BJB TC-1614 epoxy, Stycast W19 + Catalyst 9 epoxy, and West System 109 + 209 Hardener epoxy.A separate infiltration study was performed on 2.54 cm cubes whereby each cube was submerged in the infiltrant and exposed to a vacuum for durations of 1, 3, and 6 minutes. The aim of this study was to determine the time required to infiltrate each test part by using the following rate of infiltration equation [Zhou et al., 2004](cid:1872) (cid:3404)(cid:1848)(cid:3043) (cid:1850)where t is the time (s) for total infiltration, Vp is the volume of porosity (cm3), and X is the rate of infiltration (cm3/min) of the sealant into the FDM part.

The volume of porosity, Vp, was calculated by determining the difference in mass between the CAD model, or ideal test part with no porosity, and the actual test part. For the CAD model, the mass was calculated by multiplying the volume of the part (given by SolidWorks CAD software) by the density of the material, ABS-M30 (1.04 g/cm³) [Wohlers, 2011].Figure 3: Infiltration ProcedureThe mass of porosity, mp was calculated by subtracting the mass of the real part from the mass of the ideal part. Finally, the volume of porosity was calculated by using the following:(cid:1848)(cid:3043) (cid:3404)(cid:1865)(cid:3043) (cid:2025)After determining the total infiltration time, the multi-feature test part was vacuum infiltrated (Figure 3) for the calculated total infiltration.3.4 BrushingA brushing method was tested that can be implemented by consumers with limited resources in which vacuum infiltration is not possible.

Foam brushes were used over bristled brushes because they produce parts with better esthetics and minimized dimensional changes while also being able to absorb excess material where needed. Brushing was performed on the entire surface of the multi-feature test part with all 11 sealants.

Brushing was performed on the entire surface of the multi-feature test part with all 11 sealants. As noted in Figure 4, sealants can be classified as a one-part sealant (Minwax Sanding Sealer, Thompson’s WaterSeal, Polyurethane Oil Base, Polyurethane Water Base, DEFT Clear Wood Finish lacquer, IPS Weld- On 3 Cement), two-part epoxy (BJB TC-1614 epoxy, Stycast W19 + Catalyst 9 epoxy, West Marine penetrating epoxy, West System 109 + 209 Hardener epoxy), or a special case such as Hysol E-30HP which is a two-part epoxy that can be used directly with a mixing nozzle.3.2 Testing for dimensional changes and performance under fluid pressureDimensional changes of the test part were evaluated by using a SmartScope Flash 250 coordinate measuring machine (Optical Gauging Products, Inc., Rochester, NY) to measure selected features before and after sealants were applied.

A fixture (Figure 5) was fabricated to enable quick and easy registration of the parts so that accurate and repeatable measurements could be obtained. Vertical and horizontal locations along the top of the multi-feature test part were measured and include 33 total measurements.One-Part SealantTwo-Part EpoxyHysol E-30 ClFigure 4: Brushing application procedureAn aluminium-6061 test fixture (Figure 6) was manufactured using a HAAS Super Mini Mill 2 Computer Numerical Control system (Haas Automation, Inc., Oxnard, California). The fixture, along with the surface treated test part, was submerged in water and attached to a compressed air source that supplied a prescribed pressure at a rate of 6.9 kPa/min (1psi/min) and held at 276 kPa (40 psi) for at least 5 minutes. Safety concerns pertaining to a catastrophic failure due to the pressure differential were taken into account by applying the above mentioned pressure rate.

Safety concerns pertaining to a catastrophic failure due to the pressure differential were taken into account by applying the above mentioned pressure rate. A failed test part or feature was denoted by leaking and air bubble formation. Conversely, a successful test part or feature was denoted by the lack of bubble formation and a confirmation test was performed the next day to ensure the part’s sealed conditions remained unchanged.Test partFixturFigure 5: Coordinate measuring machine fixture and setupAssembly boltsO-ring SealsHigh Pressure AreaPressure source and regulatorFigure 6: Fixture for pressure testing4.0 Results4.1 Vacuum infiltrationThe 76.20 by 19.05 by 3.18 mm (3 by 0.75 by 1/8 inch) samples were cut and its cross-section was observed using a stereomicroscope as seen in Figure 7. A successful infiltration of the small test part was noted by the lack of the FDM material beads as in Figure 7 c, e, f, h, and i.

A successful infiltration of the small test part was noted by the lack of the FDM material beads as in Figure 7 c, e, f, h, and i. Conversely, unsuccessful or partial infiltration was noted when FDM material beads were clearly evident as in Figure 7 b, d, and g. Polyurethane Water based sealer was not tested for infiltration due to its freezing behavior under vacuum caused by its chemical composition. Hysol E-30CL was also not tested for infiltration due to its high viscosityabcdefghiFigure 7: Cross-sectional view of vacuum infiltrated test parts, a) No infiltration (benchmark), b) Thompson’s WaterSeal Multi-Surface Waterproofer, c) Minwax Sanding Sealer, d) DEFTClear Brushing Lacquer, e) BJB TC-1614 A/B, f) Stycast W19+Catalyst 9, g) West Marine Penetrating Epoxy, h) West System 105 Resin+209 Hardener epoxy, i) Minwax Oil Based Polyurethane (note that c, e, f, h, and i show evidence of infiltration)(10,500 cP).

Finally, IPS Weld-on 3 was not used as an infiltrant in this study due to its fast drying time (1 min), but is suggested for other application methods like dipping. Based on the results from Figure 7, the infiltration time was determined for the following candidate infiltrants: Minwax Sanding Sealer, Polyurethane Oil Based, BJB TC-1614 epoxy, Stycast W19 + Catalyst 9 epoxy, and West System 109 + 209 Hardener epoxy.Volume of porosity for the multi-feature test part was found to be 11.51 cm³, corresponding to 5.96% of the total volume (202.17 cm³). The rate of infiltration, X, varied according to the density and viscosity of each sealant. The time required to infiltrate the cubic test parts is described in Table 2.4.2 Performance under fluid pressureFigure 8 presents the performance of each sealant/infiltrant with the corresponding application method.

Pressures of 138 – 207 kPa were withstood when two coats were brushed using Polyurethane Oil Base, Polyurethane Water Base, Minwax Sanding Sealer, Stycast W19 + Catalyst 9 epoxy, or West System 105 + 209 Hardener. This pressure range was also achieved with the vacuum infiltration of BJB TC-1614. The maximum tested pressure, 276 kPa, was withstood by the multi-feature test part when brushed with two coats of BJB TC-1614. It is worth noting that the average dimensional change for brushing two coats of BJB TC-1614 was 0.231 mm.As previously mentioned, a failed test part or feature was denoted by leaking and air bubble formation. To ensure safe testing conditions, none of the parts were allowed to experience catastrophic failure (i.e., exploding). The majority of failure sites for the multi- feature test part included locations where geometry changes occurred (e.g., a flat surface transitioned to a round surface).

The majority of failure sites for the multi- feature test part included locations where geometry changes occurred (e.g., a flat surface transitioned to a round surface). Figure 9 shows the formation of bubbles at multiple locations where dissimilar surfaces intersect. Future work can include the incorporation of a fillet feature at such intersections made from a sealant such as epoxy. Since most failures occurred randomly throughout geometry changes, a definite conclusion cannot be made on whether the feature thicknesses had an influence on the failure of the parts under pressure.

The following figure shows a summary of the results obtained by the coordinate measuring machine. The average dimensional change is defined as the absolute difference between the non-treated part and the brushed/infiltrated part averaged over 5 features.

Overall, parts that were infiltrated showed less dimensional changes (average absolute change of 0.064 mm) than parts that were brushed (average absolute change of 0.073 mm and 0.139 mm when one coat or two coats of sealant were applied, respectively).typical lack of bubble formation at continuous surfacestypical bubble formation at intersection of dissimilar surfacesFigure 9: Surface-treated, ABS-M30 part under fluid pressure9 0 2 + n i s e R 5 0 1 m e t s y S t s eWy x o p er e n e d r a HO√√O0.350)0.300m m(0.250s e g n a h C0.200l a n o i s n e m Di0.1500.1000.0500.000DEFT LacquerWeld-On 3Minwax Sanding SealerMinwax Oil Base PolyurethanePRO Finisher Water-Base PolyurethaneThompson's WaterSealBJB TC-1614 A/BHysol E- 30CLStycast W19 + C9West Marine Penetrating EpoxyBrush 1 CoatBrush 2 CoatsInfiltrationFigure 10: Dimensional changes (note not all materials were vacuum infiltrated)5 ConclusionsManufacturing end-use parts using FDM has not been possible for fluid pressure applications due to part porosities, air gaps, and voids.

Improving such build defects may allow FDM technology to be used for applications where fluids are applied at low pressures including hermetic housings for biomedical devices such as pacemakers and pipes/covers for thermodynamic systems such as heat exchangers.

Vacuum infiltration and brushing were explored with the use of six sealants (DEFT Clear Brushing Lacquer, Minwax Sanding Sealer, Minwax Oil Base Polyurethane, PRO Finisher Water-Base Polyurethane, Thompson's WaterSeal Multi-Surface Waterproofer) that are readily available to consumers at hardware stores and six industrial sealants (IPS Weld-On 3 Cement, BJB TC-1614 A/B, Hysol E-30CL, Stycast W19 + Catalyst 9, West Marine Penetrating Epoxy, West System 105 Resin + 209 Hardener epoxy) that are primarily used by industrial companies for specialized repairs.In this study, post-processing a multi-feature, FDM test part (made of ABS-M30) with BJB TC-1614 had notable results: individually, brushing and vacuum infiltration allowed the test part to repeatedly hold a fluid pressure for at least 5 minutes at pressures of 276 kPa and 138 kPa, respectively.

The mean absolute difference between the non-treated part and the brushed/infiltrated part caused by the application of BJB TC-1614 was minimal (0.231 mm for brushing and 0.104 mm for infiltration). This dimensional change can be controlled through user application during brushing and vacuum infiltration. A chart (Figure 8) was developed that can be used by consumers to select appropriate application techniques corresponding to available sealant options. This study suggests FDM customers can build end-use parts for low pressure applications by applying one of the procedures explored in this research. Ease-of-use needs to be considered as well as sealant cost, drying/working times, aesthetics, as well as envinronmental and health risks when needed. Epoxies were more expensive and their work times need to be followed as compared to one-part sealants such as Minwax Sanding Sealer which is ready to apply and has a long work time.

Epoxies were more expensive and their work times need to be followed as compared to one-part sealants such as Minwax Sanding Sealer which is ready to apply and has a long work time. When precaution is taken, all sealants produced desirable aesthetic results however West Marine Penetrating Epoxy had notable accumulation during the drying process and produced unappealing surfaces. Epoxies generally had the greatest change in dimension while one-part sealants had the least with some noteworthy results as seen in Figure10. When a part is to be used rapidly, epoxies dry within 5 hours with notable sealantsWest System 105 + 209like BJB TC1614 that were able to be used after 2 hours, however other noteworthy sealants such as Minwax Sanding Sealer and polyurethanes required at least 24 hours to completely dry making them unsuitable for rapid use and quick processing.

Environmental and health risks also need to be taken into consideration as some sealants require special safety precautions during processing (e.g. fume hoods, safety equipment) as well as precautions needed for specific end- use applications (e.g. FDA approved, biomedical). It is to be noted that sealant performance and dimensional changes may vary from user to user. Precautions should be taken when using FDM- manufactured parts in fluid pressure applications as voids may give rise to catastrophic failures such as exploding. FDM consumers may apply multiple coats to seal their FDM parts, however multiple coats was not explored in depth for this study.Future WorkAlthough not captured by dimensional measurement data, sealant material tended to accumulate during drying/curing at corners, edges, crevices, and the test part’s base.

To combat sealant accumulation, a rotary arm can be used to continuously move the part as it dries, giving the final product an even and esthetically appealing surface. Employing this method to achieve evenly distributed surface coatings can allow for multiple sealing coatings while minimizing dimensional changes. Other sealants with different properties can be explored including off the shelf sealants which may be ideal to consumers with short wait times.In this study, pressure was applied gradually; however, many applications require endurance of instant pressure which should be the focus of future work. Additional work should also determine a sealant’s maximum working pressure and lifetime in various conditions such as exposure to ultraviolet light, dry/humid environments, and cyclic fluid pressure loading.

For tests involving instantaneous and high fluid pressures, special precautions should be taken since FDM-manufactured parts may fail catastrophically due to voids developed by the build process.Optimizing FDM-building variables (e.g., build orientation, thermoplastic material, raster modifications, air gaps between rasters and contours) can directly impact part sealing [Thrimurthulu et al., 2004] and have shown to directly affect mechanical properties where radiused surfaces are present [Ahn et al., 2002]. Therefore, a design of experiments including multiple factors and levels should be created that determines the influence of the various FDM- building parameters on a part’s performance under fluid pressure. Additionally, other promising sealing methods such as electroplating, vapor smoothing, and dipping in solvents may be investigated to determine their efficacy.

Additionally, other promising sealing methods such as electroplating, vapor smoothing, and dipping in solvents may be investigated to determine their efficacy. Ease-of-use needs to be documented as different methods require special processing techniques and hardware making some methods more ideal than others. Optimized building parameters should then be used concurrently with successful sealing/infiltration methods – using a single sealant or a combination of sealant and/or application methods – to fill voids within FDM-manufactured parts that will enable end-use applications in high fluid pressure environments. FDM materials other than ABS-M30 should also be tested to determine if sealing/infiltration behavior and dimensional changes vary due to the thermoplastic composition.AcknowledgementsThe research described in this paper was performed at the W.M. Keck Center for 3D Innovation at the University of Texas at El Paso (UTEP).

ALYSIS OF THE EFFECTS OF 3DP PARAMETERS ON PART FEATURE DIMENSIONAL ACCURACYTamas D. Szucs and Dermot BrabazonSchool of Mechanical and Manufacturing Engineering, Dublin City University, IrelandAbstract3D printing (3DP) is a widely investigated scaffold manufacturing process for Tissue Engineering (TE). Useful scaffold geometries should have high porosity (60-80%) with small (100-500 μm) interconnected pores. Therefore dimensional accuracy on the micron level is one of the crucial parameters of the bone scaffolds. Previously it was shown that the behavior of scaffold geometries can be well simulated with Finite Element Modeling (FEM) however the prediction of actual strength and stiffness values are dependent on dimensional accuracy. This accuracy is in turn dependent on several parameters including particle size and shape, powder- binder interaction, and machine setup.

This accuracy is in turn dependent on several parameters including particle size and shape, powder- binder interaction, and machine setup. In this work different scaffold strut sizes (0.3 - 0.5 mm) have been fabricated using two different plaster powders (zp102 and zp130) with variations in shell saturation levels, part print position, and part print orientation. The parameters for each powder were analyzed using a full 35 factorial experimental design. It was found that the part size and orientation had a significant effect on the dimensional accuracy while the influence of the shell saturation and position was relatively small.

It was found that the part size and orientation had a significant effect on the dimensional accuracy while the influence of the shell saturation and position was relatively small. The results allow for better dimensional specification for scaffold geometry fabrication by defining the process parameters in 3DP that may be used further in scaffold accuracy optimization.IntroductionTissue engineering aims to produce artificial biological substitutes by using suitable cells in combination with various materials and engineering solutions. This goal is important due to the increasing donor shortage and the potential to improve patient response with tailored and definable scaffold properties [1]. A main approach in tissue engineering is to use the patient’s own cells to build up an implant that can augment or replace a tissue function including a complete organ.

A main approach in tissue engineering is to use the patient’s own cells to build up an implant that can augment or replace a tissue function including a complete organ. The fundamental problem is that the cells can proliferate into larger colonies; however they are incapable of forming three-dimensional tissues/organs in vitro [2]. To form a three-dimensional organized tissue the cells need complex mechanical, chemical, and electrical signals which are present in their natural environment [3]. For this reason it is necessary to use a porous matrix called a scaffold to which the cells can attach, proliferate and differentiate in vitro and afterwards this can be inserted to the anatomical defect (Figure 1). This matrix is only necessary until the cells form the desired anatomical shape and gain sufficient mechanical properties to withstand the physiological loading.

This matrix is only necessary until the cells form the desired anatomical shape and gain sufficient mechanical properties to withstand the physiological loading. Therefore scaffolds are ideally made from bioresorbable materials [4].Figure 1 Scaffold use for cell proliferation to produce a functional bio-implant.In the past few years the number of investigations regarding to Solid Freeform Fabrication (SFF) techniques in manufacturing scaffolds for TE proposes have been significantly increased [5-7]. It is easy to see why this is the case as there is great potential for SFF techniques to manufacture customized products with complex internal geometries. Scaffolds need to incorporate several characteristics that are often contradictory but using FEA combined with SFF can lead to an implant that has been tailored for actual defect properties in the patient.Due to its complex role the scaffold needs to fulfill requirements in many aspects.

Appropriate micro and macro geometry including high porosity (50-80%) is required to enhance mass transport which aids nutrient and metabolic waste transport to and from the cells [8]. Moreover the pore structure has to be interconnected to promote tissue integration and vascularisation [9]. The macro geometry of the scaffold has to fit into the complex three- dimensional anatomic defect. To satisfy this requirement medical imaging technologies like Computed Tomography (CT), µCT or Magnetic Resonance Imaging (MRI) can be used for mapping the defect site [10]. Scaffolds should also possess sufficient mechanical properties that provide structural support for the cells until the native tissue builds up [2]. The scaffold is ideally made of bioresorbable material, with appropriate degradation profile that will dissolve in the body without toxic by-products, thus providing space for new tissue formation.

The scaffold is ideally made of bioresorbable material, with appropriate degradation profile that will dissolve in the body without toxic by-products, thus providing space for new tissue formation. Furthermore appropriate surface chemistry is of high importance in ensuring cell attachment, differentiation and proliferation [11]. Proceeding the SFF period numerous conventional techniques including solvent casting particulate leaching, fiber bonding, membrane lamination, and gas foaming were invented to fabricate porous structures. For example, in solvent casting/particulate leaching, pores are formed by evaporative solvent and pore-forming agent particles that are mixed and cast into a mould and leached out afterwards. This method just like other conventional techniques is incapable of precisely controlling pore size, geometry and spatial distribution. With these techniques, the internal architecture cannot be accurately designed in advance.

With these techniques, the internal architecture cannot be accurately designed in advance. Moreover there may be no guarantee of the required pore interconnectivity [2]. Using conventional techniques the size of the scaffold can be limited due to difficulties in removal of the porogens. Another important disadvantage is that some these techniques are applying organic solvents during the process that may remain in the scaffold afterwards. These are unfavorable in medical applications due to their carcinogenicity and toxicity. These techniques are also labor demanding, not repeatable and often do not control well the micro and macro geometry of the scaffold. Most of the above listed problems could be overcome by using SFF techniques [6].Probably the most significant advantage of SFF techniques over the so called conventional scaffold manufacturing techniques is that they can produce more intricate, previously designed and analyzed geometries.

Taking advantage of the possibilities of SFF we can have strict control on the porous architecture size, shape and interconnectivity. 3D printing (3DP) is one prospective SFF technique that may be used in manufacturing hard tissues. Currently 3DP is successfully used in various applications including rapid tooling and appearance modeling. However in these cases requirements are often less complex and there are several specific barriers that need to be overcome for hard tissue engineering. In order to design bone substitutes with unique, tailored mechanical properties preliminary FEA is essential. Previous studies have shown that it is possible to characterize the mechanical properties of the structure by changing the pore structure only and keeping the porosity on the same level [12].

Previous studies have shown that it is possible to characterize the mechanical properties of the structure by changing the pore structure only and keeping the porosity on the same level [12]. Furthermore, it was shown that actual geometry solutions provided from FEA cannot fully be examined experimentally due to dimensional differences between the FEA geometry and that which can be produced from the actual SFF technique [12]. In this work process parameters including powder type, shell saturation, part position, and orientation were examined with the 3D printingtechnique. This was done in order to assess the potential for this technology to meet the small scale geometric requirements of hard tissue bone scaffold structures. The results may be interpreted for later studies with biodegradable raw materials.Materials and methodsFigure 2 presents the specimen geometry that was fabricated in this work.

The results may be interpreted for later studies with biodegradable raw materials.Materials and methodsFigure 2 presents the specimen geometry that was fabricated in this work. The CAD file dimensions for the cross members were 300 x 300, 400 x 400 and 500 x 500 micrometers and 5 mm in length. Three beams with these dimensions were printed in each test sample. Screening experiments were performed by printing 0.1 to 1.2 mm square cross section beams. Printed beam cross sections of 200 x 200 micrometer and less were not stable enough after printing to support themselves. The minimum beam cross section of 300 x 300 was therefore chosen for this work. The beam sizes chosen are also typical of scaffold beam sizes used in tissue engineering. The geometries for the accuracy measurements were designed using the SolidWorks CAD (SolidWorks Inc., Concorde, MA, USA) program.

The geometries for the accuracy measurements were designed using the SolidWorks CAD (SolidWorks Inc., Concorde, MA, USA) program. A ZPrinter 310 (Z Corporation, Burlington, MA) was used to produce the 3DP structures.Row of 500 x 500 m cross sectionRow of 400 x 400 m cross sectionRow of 300 x 300 m cross sectionFigure 2 Schematic of the fabricated test frame geometry.Five different process parameters were investigated. These included location in the x and y axis on the 3DP build bed, the orientation with respect to the machine orthogonal axis, the saturation level, and the beam size (300, 400 and 500 micrometer). Three levels of each of these five parameters were chosen and used in a full factorial design of experiments.Table 1 summarizes the studied factors and the levels of these factors. For each powder, this resulted in 243 printed test frames fabricated with the associated process parameter combinations.

For each powder, this resulted in 243 printed test frames fabricated with the associated process parameter combinations. An average value and confidence interval of the beams of similar size were calculated and recorded for each frame structure. In every case the specimens were printed with 0.0875mm layer thickness, bleed compensation set to off and they were left to dry for 24 hours before depowdering. Further description of the factors is presented below.LevelsLevel 1 Level 2 Level 3Saturation level 75% 100% 125%Shell saturationActual binder/volume ratiozp 102 0.281747 0.229571 0.469578zp 130 0.172178 0.375662 0.286964Size Orientation0.3 mm 0.4 mm 0.5 mmX Y ZX positionLeft Center RightTable 1 Factors and levels in the design space for the zp102 and zp130 powders.Y positionLeft Center RightPowder typeTwo different plaster powders were used, the well known zp102 and the relatively new zp130 with the appropriate binders zb56 and zb58 respectively.

According to the manufacturer’s specification the zp130 provides improved feature definition that has been investigated in this work. The main difference in processing the two powders is that zp102 powder needs more time to reach its full strength since higher saturation levels are used for this powder both in the shell and the core. The printing speed is higher for the zp130 powder for the same reason. New batches of powder were used for this work and the particle size distributions were analyzed before being used for both powders using the Malvern Mastersizer S (Malvern Ltd., Malvern, UK) particle analyzer.Shell saturation levelThe saturation level determines the depth of binder penetration within the powder [13]. In general increasing the saturation level results in higher strength while lower saturation levels makes depowdering easier. Saturation levels can be set selectively for the shell and core of the part.

Saturation levels can be set selectively for the shell and core of the part. The default level is noted as 100% for both the shell and the core however the actual saturation is noted as binder volume ratio. This ratio indicates that on default settings the printer head uses higher saturation on the shell. The recommended saturation settings vary with the powder type. In this work saturation levels of 75%, 100%, and 125% were analyzed for both powders. These settings automatically equate to different binder/powder volume ratios depending on the powder selected in the ZPrint software. The levels were selected such that the 125% saturation of the zp130 provided the same binder/powder ratio as the 75% level for the zp102.Part positionPart accuracy may depend on the position of the part in the build chamber, although this effect may vary with machine type and actual condition. In this paper the parts were positioned to 9 different locations.

In this paper the parts were positioned to 9 different locations. The coordinates of the 9 specimens are depicted in Figure 4 and presented in Table 2. The origin used in this work is shown Figure 3.ZYY3(0, 0, 0)6X9Y584(a)7(b)Figure 3 (a) Picture of build chamber and co-ordinate system used and (b) indication of the locations of the specimens within the build chamber.21XX axisLeft (mm)Center (mm)Right (mm)Y axisLeft (mm) Center (mm) Right (mm)X 251 251 251Y 176 88 0X 125.5 125.5 125.5Y 176 88 0X 0 0 0Y 176 88 0Table 2 Coordinates of the locations at which the parts were printed.OrientationDimensional accuracy of the part is determined by the powder particle size in every direction but the accuracy in the x-y plane in more constrained by the hardware and control software. The resolution in the x (fast) and y (slow) axes depends on the printer head type (300 dpi) and the printing method.

The resolution in the x (fast) and y (slow) axes depends on the printer head type (300 dpi) and the printing method. The actual size of the features in the z direction depends on their position along the z axis. If the feature size in the z direction is smaller than the layer thickness, the replicated size will either be the layer thickness or twice the layer thickness as the feature can be positioned within one or between two layers.

This means printing a 0.4 mm thick feature with 0.0875mm layer thickness can result in five (5 x 0.875=4.375 mm) or six layers (6 x 0.875= 5.25 mm) according to its position along the z axisThe accuracy of the 3DP system in the x, y, and z directions was assessed by printingspecimens with the examined dimensions in these directions as shown in Figure 5.ZYXFigure 5 Schematic of the orientation of the printed test frames in location three, six and nineMeasuring AccuracyThe beams were measured using a Reichert MeF2 (Reichert, Vienna, Austria) inverted light optical microscope from one direction, that is the beams that were oriented in the x and y directions were measured in the x-y plane while the beams in the z direction were measured in the y-z plane. First, the captured images were thresholded and then the average chord length in the horizontal direction was calculated using the Buehler Omnimet image analysis software (Buehler Ltd. Lake Bluff, IL, USA).

Lake Bluff, IL, USA). On each specimen the nine beams were measured separately, the average of the three measurements for beams of similar size was calculated and recorded. An accuracy value of comparison for the beams was calculated by dividing the measured values by the nominal dimensions. The results of the full factorial design were analyzed and evaluated with Design-Expert 7.0 design of experiments software (Stat-Ease Inc. Minneapolis, MN, USA).Results and discussionThe results for the particle size distribution measurement of both powders are shown in Figure 6. There was no significant difference between the mean particle size diameters: it was 41.03 m for the zp102 and 40.79 m for the zp130. The distribution curves showed a noted difference between 2 and 10 m with a reduced amount of particles present in the zp130 powder. The zp102 had fewer particles between 10 µm and 40 µm but more particles below 10µm.

The zp102 had fewer particles between 10 µm and 40 µm but more particles below 10µm. For the zp102 19.3% of the powder particles were between 1 and 10 µm and 32.6% were between 10 and 40 µm. For the zp130 powder these figures were 11.2% and 43.3% respectively.600.1090 10010.020Particle Diameter (µm.)Volume (%)0 101.0305080100.070101000.04010.090 1001.0100.01000.040302010708050600Particle Diameter (µm.)Volume (%)0 100.10.11.010.0100.01000.0 0.1 1.010.0 100.0 1000.0(a)(b)Figure 6 Particle size distributions for (a) zp102 and (b) zp130 powders with the x-axis in micrometers and the secondary y axis shown in percentage.To find the factors that have a significant influence on the accuracy analysis of variance (ANOVA) was carried out.

After eliminating the insignificant interactions the ANOVA results table for the two powders are presented separately in Table 3 and Table 4.SourceModel A-Saturation B-Size C-Orientation D-X position E-Y position AB AD CD A^2 B^2 Residual TotalSum of squares Degrees of freedom Mean square 44363.55 12 1 2.23 123249.45 1 194270.12 2 1 183.27 2701.25 1 1269.73 1 1997.97 1 3030.85 2 4589.95 1 3766.78 1 263.19 230 - 242532362.57 2.23 123249.45 388540.24 183.27 2701.25 1269.73 1997.97 6061.70 4589.95 3766.78 60533.63 592896.20F value168.56 0.01 468.29 738.14 0.70 10.26 4.82 7.59 11.52 17.44 14.31 - -P value < 0.0001 0.9268 < 0.0001 < 0.0001 0.4049 0.0015 0.0291 0.0063 < 0.0001 < 0.0001 0.0002-Table 3 ANOVA table for zp102SourceModel A-Saturation B-Size C-Orientation D-X position E-Y position AB AC BC CD CE B^2 D^2 E^2 Residual TotalSum of squares Degrees of freedom Mean square 48212.41 18 6606.40 1 155766.06 1 333021.54 2 34.15 1 1 471.62 1085.68 1 5947.42 2 3979.49 2 870.36 2 957.72 2 13058.95 1 540.03 1 707.35 1 154.85 224 - 242867823.29 6606.40 155766.06 666043.07 34.15 471.62 1085.68 11894.83 7958.98 1740.73 1915.44 13058.95 540.03 707.35 34686.67 902509.97F value311.35 42.66 1005.91 2150.59 0.22 3.05 7.01 38.41 25.70 5.62 6.18 84.33 3.49 4.57-Table 4 ANOVA table for zp130From the results it can be concluded that the F value was high for both models which means the model was significant and there was only a 0.01% chance that this occurred due to noise.

The F values for the factors were calculated by dividing the means square of the factor by the residual mean square. Results with p values less than 0.05 indicate that they are significant factors. According to this, significant factors for zp102 are size, orientation, and y position and for zp130 saturation, size, and orientation are. Several other model terms, interactions are also concerned significant. The contribution of the most significant factors in is presented at Figure 7(a)(b)Figure 7 Contribution of the factors in the model for (a) zp102 and (b) zp130Figure 8 represents the effect of the beam size and the saturation on the normalized accuracy which is noted as the percentage of measured divided by nominal on the y-axis of these graphs. This represents the percentage increase of the prototype dimension over the dimension set in the CAD file. It can be concluded that increasing the size considerably increased the accuracy for both powders.

It can be concluded that increasing the size considerably increased the accuracy for both powders. There was little difference between the zp102 and zp130 powder for the 0.3 mm beam size (see Figure 8 (a) and 8 (b)). The accuracy was 20% higher with the zp130 powder compared to zp102 powder for the 0.5 mm beam. According to the results of ANOVA in Table 3P value < 0.0001 < 0.0001 < 0.0001 < 0.0001 0.6391 0.0823 0.0087 < 0.0001 < 0.0001 0.0042 0.0024 < 0.0001 0.0631 0.0337-saturation factor was not significant for the zp102.

Using the zp130 powder the higher saturation improved the accuracy although its influence was very small compared to beam size.(a)(b)Figure 8 Size and saturation factor influence on accuracy for (a) zp102 and (b) zp130The orientation had the most powerful influence in both powders and especially so for thezp130, see Figure 9 (a) and (b).The results that are presented in Figure 9 indicate that there was no considerable difference in the x and y direction for the zp102 but there was 30% (measured/nominal) difference for the zp130. In the z direction the manufactured beam size was significantly lower and closer to the set size. However there was a very large difference between the x and y, and the z axis.

In the z direction the manufactured beam size was significantly lower and closer to the set size. However there was a very large difference between the x and y, and the z axis. This is probably due to the fact that manufacturing in the z direction is less sensitive to the printer head positioning which may be responsible for the higher inaccuracy in the x-y plane.Measured / Nominal [%]Measured / Nominal [%](a)(b)Figure 9 Percentage beam accuracy in relation to beam size and orientation at a saturation of 100% for (a) zp102 and (b) zp130Figure 10 shows the actual sizes of the beams as measured with the Reichert microscope. Examining the actual average measured values for the sizes of 0.3 and 0.4 mm beams it can be established that there is no difference between the figures.

Examining the actual average measured values for the sizes of 0.3 and 0.4 mm beams it can be established that there is no difference between the figures. The difference between the width of the 0.5 mm and the 0.3 - 0.4 mm beams was between 90 and 120 μm indicating that the machine used the same number of layers when it printed the 0.3 and 0.4 mm size beams and laid one extra layer to print the 0.5 mm beams. These results confirm that the accuracy of the beams in the z direction depends on their actual position along the z axis as described above.Figure 10 Actual measured beam sizes in the z direction at 100% saturation in position 5Comparing the results for the two powders where the same binder/volume ratio was used (that is 75% for the zp102 and 125% for the zp130) no significant difference in accuracy was detected between the use of the two powders.

The saturation settings parameter had very little influence on the accuracy in case of the zp102 powder (see Figure 9 (a) relative to Figure 11 (a)). The effect of saturation was however measurable for the zp130 (see Figure 9 (b) relative to Figure 11 (b).(a)(b)Figure 11 Percentage beam accuracy in relation to beam size and orientation at a saturation of 75% for (a) zp102 and at a saturation of 125% for (b) zp130According to the ANOVA tables the x position is not a significant factor but the part position along the y axis may have some contribution in the final part accuracy. The accuracy maps of the 3D printer were drawn up from the results of the nine different locations, see Figure 12 and 13. The positions of best accuracy are indicated with “+” symbol and the locations that provided the least accurate beams are noted with “–” symbol.

The positions of best accuracy are indicated with “+” symbol and the locations that provided the least accurate beams are noted with “–” symbol. The results did not show significant differences; however with the zp102 powder the positions closer to the x axis did show better accuracy.X(a)(b)(c)Figure 12 Accuracy map of the build chamber for zp102 in (a) x, (b) y, and (c) z orientations.X(a)(b)(c)Figure 13 Accuracy map of the build chamber for zp130 in x (a), y (b), and z (c) orientations.Conclusions and further directionsSaturation was not seen to have a significant effect for the zp102 powder but larger saturations were seen to increase the accuracy slightly for the zp130 powder. Changing the saturation showed very small influence on the accuracy but low shell saturation levels made the samples very fragile that rendered the specimen removal and depowdering difficult.

Changing the saturation showed very small influence on the accuracy but low shell saturation levels made the samples very fragile that rendered the specimen removal and depowdering difficult. Comparing the two powder types it can be concluded that there is no substantial difference in accuracy in the examined size range.The sizes produced in the x and y direction were approximately 2.5 times their nominal dimensions and in the z direction were approximately 1.5 times their nominal dimensions. AnYYimprovement in accuracy was noted for larger beam sizes. The measurements in the x and y orientations showed were similar for the zp102 but were better in the x-direction compared to the y-direction for the zp130. The closest recorded result to the 300 m size was 464 m which was obtained at 75% saturation, in the z-direction, and with the zp130 powder.

The closest recorded result to the 300 m size was 464 m which was obtained at 75% saturation, in the z-direction, and with the zp130 powder. The zp130 powder, in the z-direction and with saturation at 75% and 100% produced the smallest deviations from the actual beam set dimension. The lowest percentage deviation was noted for the 0.5 mm with the zp130 powder at about 30%. The zp130 powder produced a deviation on average 51 m better than with the zp102 in the z-direction. The much better accuracy in the z-direction indicates that the positioning of the printer head in the x and y direction may be the main boundary in manufacturing small features.Two important factors however affecting the accuracy at this scale are the number of printer head nozzles that are firing and the actual size of the powder layer brought across during the process. The printer head nozzles are arranged in a zigzag pattern with diameters of 60 m with approximately 25 m spacing between them.

The printer head nozzles are arranged in a zigzag pattern with diameters of 60 m with approximately 25 m spacing between them. The firing of three of these nozzles would equate to a line of binder of approximately 255 m and four to a line of 340 m. The discrete nature of this process results in a compromised resolution in the y direction on levels examined in this work. The powder layer will also be variable to the extent that the powder with particle sizes ranging from 0 to 200 m will lead to a variable layer height somewhere in between depending on packing density and pressure that may also contribute to the accuracy in the z direction. The position along the z axis may also influence the accuracy however this effect can be easily accounted for as described above.

The position along the z axis may also influence the accuracy however this effect can be easily accounted for as described above. Not withstanding these complications binder bleeding also presents a considerable possibility for deterioration in dimensional tolerances which could justify the oversized beams noted in this work.The printed 500 μm beam size is viable for scaffold geometries although the accuracy especially in the x-y plane is not acceptable for tissue engineering application. Part position in the build chamber changes the accuracy somewhat but significant trends could not be detected and its contribution in the results was negligible compared to the other factors. Further investigations will include various biocompatible powders including starch, and biodegradable polymers with appropriate binders using various particle sizes and particle size distributions.

Further investigations will include various biocompatible powders including starch, and biodegradable polymers with appropriate binders using various particle sizes and particle size distributions. Sieving the powder is one way to make batches with different particle sizes although this process for large amounts is found to be inefficient. The effect of the saturation settings on mechanical properties will be further investigated in the future.Acknowledgements This research has been supported by a Marie Curie Early Stage Research Training Fellowship of the European Community's Sixth Framework Programme under contract number MEST-CT- 2005-020621.1. 2.3.4.5.6. 7.8.9. 10. 11.12.13.ReferencesHutmacher, D.W., Journal of Biomaterials,VSP, 2000, 21, pp. 2529-2543. Sachlos, E., and Czernuszka, J.T., European Cells and Materials, European Cells & Materials Ltd, 2003, 5, pp. 29-40.

ALYSIS OF THE INFLUENCE OF VISCOELASTICITY IN CURL DEVELOPMENT IN SLSNM Jamal & KW DalgarnoSchool of Mechanical Engineering, University of Leeds, Leeds, LS2 9JT, UK.AbstractCurl in selectively laser sintered parts arises mainly from thermal distortion of parts within the build volume during processing. This results in nominally flat surfaces which lie horizontally in the part bed becoming warped. This paper reports on the use of finite element techniques to model curl development in polymer materials, and in particular on the influence of viscoelasticity on how curl develops. The development of time-dependent material models is reported, and the results of the implementation of these models presented, together with a comparison of the results with experimental data.IntroductionIn common with most rapid prototyping processes, SLS does not produce parts as accurately as more conventional manufacturing processes, with accuracies of ± 0.2 mm commonly quoted [1].

Inaccuracy in SLS can arise from (i) data exchange, (ii) the inherent mis-match between a 3D object and a collection of 2D layers, (iii) the resolution of the laser and laser positioning system, and (iv) shrinkage and warping as a result of the distributions of thermal energy and melting/solidification mechanisms [2]. Choi and Samavedam [2] have noted that it is the last of these sources of inaccuracy which are currently least well understood, and present methods by which sources (i) to (iii) can be minimised.One result of shrinkage and warping in SLS is the phenomena known as curl, which results in nominally flat surfaces which lie horizontally in the part bed becoming warped. Figure 1 illustrates this schematically. Previous work on the selective laser sintering of polycarbonate [3] has identified two major regions of curl within a part.

Figure 1 illustrates this schematically. Previous work on the selective laser sintering of polycarbonate [3] has identified two major regions of curl within a part. The first of these appears over approximately the 2.5 mm of part closest to the edge of the part and on parts of with the geometry shown in Figure 1 has a radius (RI in Figure 1) of around 20 mm. The second extends from the end of RI approximately 10 mm further towards the centre of the part and has a radius (RII in Figure 1) of around 60 mm. RI curl has been seen to exist from the sintering of the first layer in part build, while RII curl develops as layers are added to the part.

RI curl has been seen to exist from the sintering of the first layer in part build, while RII curl develops as layers are added to the part. RI curl is considered to arise from densification and shrinkage effects during processing of the first layer, whereas RII curl is believed to arise from thermally induced stresses which develop as subsequent layers cool onto layers which have already been processed [4].Previous work [5] attempted to model RII curl development in the SLS of polycarbonate using finite element methods, and did provide some insight into the mechanism of curl development. However, within these models assumptions were made regarding the material properties, and, in particular, no account was taken of the effects of viscoelasticity.

However, within these models assumptions were made regarding the material properties, and, in particular, no account was taken of the effects of viscoelasticity. The aim of the work described in this paper was to develop improved finite element models in order to provide greater understanding of curl development.349Figure 1 - Effect of Curl on Component Geometry of a 90 x 25 x 7 mm PartApproach to AnalysisAnalysis of curl development in SLS requires that a coupled thermal-stress analysis be used. Such analyses can be either sequentially coupled, where the results of a heat transfer analysis is used to provide temperature data to a stress analysis, or fully coupled, where the heat transfer and stress analyses are solved simultaneously.

A sequentially coupled analysis may be performed where it can be assumed that the thermal analysis would not be affected by the results of the stress analysis, and this assumption has been made in the work reported here.The models were based on the manufacture of a simple test part on a DTM Sinterstation 2000, in which individual layers were 90 x 25 mm in area, and ten layers, of nominal thickness 0.127 mm, made up the part. The laser power, laser speed, and scan spacing were 11 W, 1188.87 mm/s, and 0.25 mm respectively. The f.e. code used was ABAQUS. This code allows for elements to be activated and deactivated through the course of an analysis, which was used to simulate the addition of material, and also has a wide range of material models, which were of value in modelling viscoelasticity.Heat Transfer Analysis ResultsThe heat transfer modelling is reported in full elsewhere [6].

As the main purpose of this paper is to examine the effects of viscoelasticity, emphasis will be placed on the stress analysis and only relevant results from the heat transfer analysis are presented. The heat transfer analysis considered both the part itself and the powder bed, and Figure 2 shows the temperature results from the heat transfer model for a typical layer deposition/heating/lasing/cooling sequence, and for final cooldown. In Figure 2 (a) the points in time at which the elements making up a particular layer were activated are illustrated, as is the instant at which sintering was modelled.Stress Model - GeometryFigure 3 shows the f.e. mesh used in the stress analysis. As for the thermal analysis a 2D representation has been used, with advantage taken of symmetry to half the size of the model.

mesh used in the stress analysis. As for the thermal analysis a 2D representation has been used, with advantage taken of symmetry to half the size of the model. Within the stress analysis the powder bed was not considered to have any significant influence on the distortion of the manufactured part, and so the stress analysis considered only the manufactured part itself. In modelling the powder bed it has been assumed that the most significant effect the powder bed would have would be in restricting the contraction and expansion of the layers in the x direction, with the powder expected to have little influence on the behaviour of the part in the z direction. This led to the use of a rigid surface to represent the powder bed, as this allows frictional behaviour between the part and powder bed to be specified,350whilst the rigid assumption removes the need for any interaction between part and powder bed in the z direction.

It can also be seen in Figure 3 that the four rows of elements are used to make up layer 1, with one row of elements for each of the subsequent layers: this reflects the effect of bonus-z, and the layer thicknesses used are based on experimental measurement of the thickness of parts with the same section, but comprising 1 to10 layers [6]. The nodes forming the elements shown in Figure 3 were augmented by a set of “dummy” nodes between elements which allowed elements, when activated, to be activated in a stress free state, following standard modelling practice [7].

The density of the part was measured to be 497 kg/m3, and the coefficient of friction between the bottom surface of the part and the powder bed was defined as 0.38, based on experimental measurement at room temperature [6]. The development of an appropriate value for the coefficient of thermal expansion proved difficult. A standard value for polycarbonate would be 6.5 x 10-5 /°C, however this value had previously been found [5] to grossly overestimate curl, with a value of 2 x 10-6 /°C more representative of the behaviour observed in SLS, and this value was also adopted in the work reported here. This value was applied to all elements except the lower two rows shown in Figure 3. The thermal model indicated that the average nodal temperatures in these elements did not exceed the glass transition temperature (175°C). This, allied to observation by Ho et. al.

The thermal model indicated that the average nodal temperatures in these elements did not exceed the glass transition temperature (175°C). This, allied to observation by Ho et. al. [8] that this area of layer 1, which in effect represents part of the bonus-z, was only lightly fused, led to the assumption that the expansion coefficient would be very low in this region, and so the expansion coefficient for these two rows of elements was set to zero. To develop values of the instantaneous modulus tensile tests at a range of temperatures were carried out, and the results of these can be seen in Table 1. The f.e. code used linear interpolation to develop modulus values throughout the temperature range. Modulus values above a temperature of 150°C could not be obtained, and it was assumed that above this temperature the modulus value measured at 150°C applied. A Poisson’s ratio of 0.385 was used, based on room temperature measurements reported by Bushko and Stokes [9].

A Poisson’s ratio of 0.385 was used, based on room temperature measurements reported by Bushko and Stokes [9]. The time dependent response of the polymer was defined in the stress model351through a viscoelastic material model which forms part of the ABAQUS code. This initially simulates the time dependent response at a reference temperature based on stress relaxation data provided at that temperature, and then employs the WLF approximation to predict the time dependent response at all other required temperatures [10].Axis of SymmetryMagnified belowTen-layer partRigid surfacezRigid body reference nodex9 layers (cid:1) 0.127Layer 10 Layer 9 Layer 8 Layer 7 Layer 6 Layer 5 Layer 4 Layer 3 Layer 20.1270.13 0.15Layer 10.1830.3125Figure 3 - Stress Analysis Mesh.

Dimensions in mm.Temperature (ºC) Modulus (MPa)20 329.3120 221.4135 204140 106145 61147 38.4150 11.8Table 1 - Measured Instantaneous Modulus ValuesThe reference temperature behaviour was modelled using the Prony series, which is based onn spring/dashpot Maxwell elements in parallel [11] and can be expressed as,)t(E Eon(cid:2)(cid:3) (cid:1) 1i(cid:1) 1E Eio(cid:1) 1(cid:2)exp((cid:2)/t(cid:1)im(cid:2))where E(t) is the time dependent modulus, E0 is the instantaneous modulus, Ei is the elasticity of the spring in the ith Maxwell element, and (cid:1)mi is the relaxation time constant of that element, and t is time. The implementation of the material model in ABAQUS requires an instantaneous modulus value together with values of E(t). The reference temperature employed in the analysis was 145°C, and so the instantaneous modulus at that temperature from Table 1 was used.

The reference temperature employed in the analysis was 145°C, and so the instantaneous modulus at that temperature from Table 1 was used. Figure 4 (a) shows the results of a stress relaxation test – a table of values from that figure were defined in the model, and from this information the ABAQUS programme calculated the n, Ei, and (cid:1)mi values which were used in the analysis.352The WLF approximation is based on the observation that E(t) curves for a particular polymer developed at different temperatures and plotted on a log time scale are similar in shape but horizontally shifted along the log time scale. This is illustrated in Figure 4 (b) for three temperatures: Ts, T1, and T2. This means that if the modulus-log time curve is known at a reference temperature, for example Ts, the curves at other temperatures can be estimated by applying a shift of log(aT) on the log time scale [12].

The WLF approximation is commonly expressed as,(cid:1)(cid:1) (cid:1)T(C 1 (cid:2) T()T s )T s(cid:3)log()a TC 2 where C1 and C2 are constants which must be derived for a particular reference temperature, and T is the temperature at which the modulus-log time curve is required. The WLF approximation is generally held to be accurate over the range Ts±50°C [11]. The reference temperature used in this study was 145°C, and from stress relaxation tests values for C1 and C2 of 16.47 and 115.549°C respectively were used.

The reference temperature used in this study was 145°C, and from stress relaxation tests values for C1 and C2 of 16.47 and 115.549°C respectively were used. The development of these values in full is outwith the scope of this paper, a full treatment of their development may be found in [6].(a)70(b)60E(t)50) a P M(40TsT1T2s u u d o Ml30log aT1log aT22010001020304050log (time)Time (sec)Figure 4 - (a) Stress Relaxation Test on SLS'd Polycarbonate Sample at 145ºC, and (b) Creep or Stress Relaxation Curves Obtained at Different Temperatures (Ts < T1 < T2), after [12].Stress Analysis - Loading and Boundary ConditionsThe only boundary conditions applied were to enforce symmetry on the nodes at the left hand side of the mesh, and to prevent the rigid body from movement.

Two forms of loading were applied, gravity loads and thermal loads, to two types of analysis step, powder addition and sintered layer cooling.Powder addition was modelled by adding a uniform pressure representing the weight of the added powder to then top of the active mesh, and by reading the temperature data from the heat transfer analysis through to allow the thermal strains to develop. In sintered layer cooling steps the elements forming the most recently added layer were activated, given their weight, and the uniform pressure applied in the previous step to represent their weight removed. The temperature data was also read from the heat transfer analysis, however, temperature data was only read through once the temperature had reduced below 175°C. Dynamic Mechanical Analysis testing with the polycarbonate material identified the glass transition temperature for the material to be 175°C.

Dynamic Mechanical Analysis testing with the polycarbonate material identified the glass transition temperature for the material to be 175°C. It was considered that the polymer would not sustain stresses at a353temperature above the glass transition temperature, and so the data from above that temperature was discarded, as shown in Figure 2 (a). In addition the values of the expansion coefficient for any part of the mesh where the temperature exceeded 175°C were set to zero.Stress Analysis - ResultsFigure 5 shows the predicted profiles of the upper and lower surfaces of the part at the end of the analysis (at room temperature). Table 2 indicates how the maximum vertical distance of the lower surface of the part changes between analysis steps within the analysis as a whole, although for brevity only the changes associated with modelling four of the layers are listed.

These results show that, the addition of the colder powder causes an increase in the predicted curl, the laser sintering step causes a decrease in the predicted curl, and that as the new layer cools, the predicted curl increases again. The results also show that these effects generally reduce in scale as the number of layers increases, as the amount of material available to resist the distortion increases. Table 2 also shows that the influence of the final cooldown step on the final predicted curl is high.)m m2.2Axis of symmetry( e c a f r u s d g i ri21.81.6m o r f1.41.2n o i t c e r i d - z10.8n0.6ie c n a t s Di0.40.200102030405060708090Distance in x-direction from edge of part (mm)Figure 5 - ◊ Experimental Measurement of Part Lower Surface, x Experimental Measurement of Part Upper Surface.

Lines Show Profiles Predicted by Analysis.Experimental Measurement of CurlComponents with the geometry described above were manufactured using the processing conditions also described above. These were then cleaned to remove any residual powder, and the upper and lower profiles of the components measured by traversing the components under a non-contacting laser displacement meter using a flat table driven by a spring loaded micrometer. Typical results are shown in Figure 5.

Typical results are shown in Figure 5. The results appear to exhibit scatter: in fact the apparent scatter is of the order of the powder particle size, and is considered to be a true representation of the outer surface of the components.354DiscussionCurl of the type investigated here is considered to arise in two ways - cooldown curl which generates a bulk deformation of a component as it cools from the operating temperature to room temperature, and in-build curl which only affects the lower of a part as the upper surface is continually levelled as each new layer is added. Figure 5 shows clearly that the model overestimates the cooldown curl, which, excepting the edge effect, is very small.

It is believed that this is as a result of the modelling assumptions made regarding the lower two layers of elements having a zero coefficient of expansion - if these two layers of elements are given the same coefficient of expansion as the other layers then the modelled cooldown curl decreases to be of a similar order of magnitude as that observed experimentally.Stage in AnalysisMaximum distance of bottom surface of build from the rigid surface (mm)Layer 1 Cooling Layer 2 Powder Addition and Heating Layer 2 in melt state following sintering Layer 2 Cooling Layer 3 Powder Addition and Heating Layer 3 Laser Sintering Layer 3 Cooling Layer 9 Powder Addition and Heating Layer 9 Laser Sintering Layer 9 Cooling Layer 10 Powder Addition and Heating Layer 10 Laser Sintering Layer 10 CoolingViscoelasticity ConsideredStart of stage 0 1.73 x 10-4 8.97 x 10-4 8.06 x 10-6 1.65 x 10-4 1.11 x 10-3 2.22 x 10-5 3.58 x 10-4 7.4 x 10-4 2.15 x 10-4 4.03 x 10-4 7.81 x 10-4 2.82 x 10-4End of stage 1.73 x 10-4 8.97 x 10-4 8.06 x 10-6 1.65 x 10-4 1.11 x 10-3 2.22 x 10-5 1.86 x 10-4 7.4 x 10-4 2.15 x 10-4 4.03 x 10-4 7.81 x 10-4 2.82 x 10-4 0.1542Viscoelasticity not considered End of stage 2.45 x 10-3 6.55 x 10-3 4.95 x 10-5 2.21 x 10-3 1.39 x 10-2 5.78 x 10-5 2.23 x 10-3 9.65 x 10-3 7.17 x 10-4 3.67 x 10-3 9.23 x 10-3 9.44 x 10-4 0.241Table 2 - Changes in Predicted Curl Through AnalysisThis means that the apparent good correlation between analysis and measurement shown in Figure 5 for the lower surface of the part is false - if the overestimated cooldown curl is removed then the analysis significantly underestimates the curl on the lower surface.

This is also most likely related to the coefficient of expansion - a very low value for this was chosen on the basis of previous work with simpler material models. It is considered that the more representative material behaviour within the analysis presented here has made that assumption obselete. Further work will examine this in more detail.The stress model with minor modifications was implemented to investigate the effect of viscoelasticity on curl. The modified model did not take into account the viscoelastic effect of the polymer, i.e. the modulus was time-independent but remained temperature dependent as specified in Table 1. Other definitions in the modified model were identical to those mentioned in the original viscoelastic stress model. Table 2 shows how the maximum distance of the bottom surface of the build, with viscoelasticity considered, compares with corresponding distances with viscoelasticity not considered.

Table 2 shows how the maximum distance of the bottom surface of the build, with viscoelasticity considered, compares with corresponding distances with viscoelasticity not considered. The general observation seen is that SLS operations suggested by the viscoelastic model to cause rise and reduction in curl, mentioned355above, have an identical effect if viscoelasticity is not considered. However, the results in Table 2 suggest that viscoelasticity does play a significant role in the development of curl.ConclusionThis paper has presented an analysis of curl development in the SLS of polycarbonate which includes the definition of a full description of the viscoelastic response of the polymer. The analysis gives a qualitative insight into the development of curl, and of the effect of viscoelasticity in that, but further work is required to develop quantitative agreement with experimental observations.References1. Shellabear, M.

Selective laser sintering of an amorphous polymer - simulations and experiments. Proceedings of the IMechE Part B, 1999, Vol. 213, pp 333-349.5. Dalgarno, K.W., Childs, T.H.C., Rowntree, I. and Rothwell L., Finite element analysis of curl development in the selective laser sintering process, In Solid Freeform Fabrication Symposium 1996, proceedings of the conference held in Austin, Texas, August 12-14, 1996, ISSN 1053-2153, pp. 559-566 (The University of Texas at Austin).6. Jamal, N.M. Finite element analysis of curl development in the selective laser sintering process. PhD Thesis, University of Leeds, 2001.7. ABAQUS/Standard Example Problems Manual Version 5.5, Vol. 1, 1995, Hibbitt, Karlsson and Sorensen, Inc.8. Ho, H., Gibson, I. and Cheung W. Effects of Energy Density on Bonus Z, Surface Roughness and Warpage of Selective Laser Sintered Polycarbonate, Proceedings of the Eighth International Conference on Rapid Prototyping, held in Tokyo, Japan, June 12-13, 2000, pp.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperANALYSIS OF THE SHIELDING GAS DEPENDENT L-PBF PROCESS STABILITY BY MEANS OF SCHLIEREN AND SHADOWGRAPH TECHNIQUESB. Hoppe1, S. Enk2, J. H. Schleifenbaum2,31SLM Solutions Group AG, 23560 Lübeck, Germany 2Fraunhofer Institute for Laser Technology, 52074 Aachen, Germany 3Chair for Digital Additive Production DAP, RWTH Aachen University, 52074 Aachen, GermanyAbstractEnsuring a robust and reproducible Laser Powder Bed Fusion (L-PBF) process depends on the design of the shielding gas flow in the corresponding manufacturing system. The gas flow assures the removal of particles from the process zone that emerge from the interaction with the laser irradiation. Minor disturbances may influence the stability of the L-PBF-Process and cause defects in final parts produced.

Minor disturbances may influence the stability of the L-PBF-Process and cause defects in final parts produced. The objective of this work comprises visualizing the convection processes and particle dynamics to analyze their influence on process stability. Therefore, a high speed camera based Schlieren and Shadowgraph setup is used to visualize convection flows as well as trajectories of metallic condensate and particles with up to 10,568 fps. This arrangement allowed investigating the influence of shielding gas flow conditions on single melt tracks.

This arrangement allowed investigating the influence of shielding gas flow conditions on single melt tracks. Corresponding results and studies on the interaction between laser irradiation and particles in varying shielding gas flow conditions are contents of this work.Keywords: Laser Powder Bed Fusion; Schlieren; Shadowgraph; Process Development; Shielding gas flow; Process stability; 316L stainless steelIntroductionThe shielding gas flow becomes ever more important in modern L-PBF machines with increasing build volumes and numbers of simultaneously working laser sources. Therefore, it is critical to understand how the shielding gas flow influences the stability of the Laser Powder Bed Fusion process. It is known that material defects resulting from an insufficiently adjusted process gas flow might affect the tensile strength of parts produced as shown by Ferrar et al. [1]. Their work suggests, that differing shielding gas flow conditions cause the defects.

[1]. Their work suggests, that differing shielding gas flow conditions cause the defects. New imaging approaches can help to characterize different shielding gas flow mechanisms like the gas convection that might lead to an unstable process. A high speed Schlieren and Shadowgraph system is suitable for visualizing convection of hot gas and particle movement during the exposure of the metal powder bed. The Schlieren as well as the Shadowgraph method enables the visualization of convection processes in transparent media. With the Schlieren method, the local fluctuation of the refractive index is exploited by inhomogeneities of density in the medium under consideration. Parts of the refracted light can be prevented from reaching a screen by means of a cutoff. The shape that emerges on the screen is called a Schliere (german for "smear"). The shadow method uses the shadow cast by objects.

The shape that emerges on the screen is called a Schliere (german for "smear"). The shadow method uses the shadow cast by objects. Since this method does not require a cutoff, the luminous efficacy is much higher than with the Schlieren method. The high temporal resolution of the high speed camera enables researchers to see small changes in the range of microseconds that would otherwise be overseen.1523Bidare et al. [2,3] have shown that it is possible to visualize the convection of argon gas entraining metal vapor while being ejected from the melt pool. This was enabled through a custom Schlieren setup and a high speed camera. They have shown that the melt pool geometry is decisive for the particle ejection behavior. Different variations of laser power PL and scanning speed vscan lead to changing ejection behaviors. Hoppe et al. were able to verify these findings with a similar setup and widened the scope of analyzed parameters [4].

Hoppe et al. were able to verify these findings with a similar setup and widened the scope of analyzed parameters [4]. Metal condensate is generally entrained in the convection flow of heated shielding gas. The appliance of Shadowgraph imagery in addition to the Schlieren images allowed for clearly linking the condensate movement to the convection flow. This holds true when applying a forced shielding gas flow to the process. Interaction between the metal condensate and laser beam was observed. Furthermore, single melt tracks were analyzed via particle tracking to measure the particle ejection behavior at varying laser power PL and scanning speed vscan. The ejection angle of particles changes according to the convection flow of the heated shielding gas. Increasing PL or decreasing vscan lead to a higher ejection angle. The effect of adjusting parameter values for PL is observed to be more dominant than similar modifications for vscan.

Increasing PL or decreasing vscan lead to a higher ejection angle. The effect of adjusting parameter values for PL is observed to be more dominant than similar modifications for vscan. However, it was not analyzed in detail how this influences part quality.Hoppe et al. focused solely on the observations made by Schlieren and Shadowgraph. This work will therefore focus on further analyzing the interaction of laser beam and ejected particles from the melt pool as well as the resulting part topology. Varying shielding gas flow conditions and their influence on the L-PBF process stability will be looked at in more detail.Experimental Setup and MethodologyExposing two scanning patterns with a laser beam allows analyzing the influence of the shielding gas flow on the process stability. A high speed camera captures Schlieren and Shadowgraph images during the exposure process for visualizing the convection flows of the shielding gas and the particle ejection behavior.

A high speed camera captures Schlieren and Shadowgraph images during the exposure process for visualizing the convection flows of the shielding gas and the particle ejection behavior. The surface topology and melt track geometry are evaluated by means of reflected light microscopy and SEM images. Correlating these results to the Schlieren and Shadowgraph images should than provide explanations on the mechanisms leading to an unstable melting process. The employed experimental setup and analyzing methods are described in the following.The experimental setup is based on a modified AconityMINI machine from Aconity3D GmbH for having the opportunity of controlling and creating a desired process gas atmosphere. The system has openings at the opposing ends of the process chamber with 4 mm thick glass windows installed therein. This enables the light to pass unhindered through the chamber during the L-PBF process in order to run the Schlieren system.

This enables the light to pass unhindered through the chamber during the L-PBF process in order to run the Schlieren system. Additionally, the powder coater has an aperture so that it can remain in the process chamber during the process and allow measurements to be made. The setup corresponds to the setup used and described by Hoppe et al. in [4]The z arrangement described by Settles [5] is used for the mirror Schlieren system in this work, but has been customized so that it is compatible with the described machine type. The Schlieren system consists of an illumination side and an imaging side, which are arranged point- symmetrically to the process chamber center.A mounted LED MCWHL5 from Thorlabs produces the cold white (6500 K) light required for the Schlieren imagery at a minimum light output of 800 mW. This light is focused by a condenser lens onto a variable slit diaphragm and then reflected via a deflecting mirror onto a first1524off-axis parabolic mirror (OAP).

This light is focused by a condenser lens onto a variable slit diaphragm and then reflected via a deflecting mirror onto a first1524off-axis parabolic mirror (OAP). The reflection at the OAP leads to a collimation of the light, which enters the process chamber through the first glass window. The light illuminates the observation plane which is approx. 230 mm away from the second glass window. The light leaves the process chamber through the second glass window and is refocused by a second OAP onto a cutoff (consisting of a standard razor blade) via a second deflecting mirror. The cutoff is positioned in the focal point of the OAP and set to reflect 75% of the focused light.

The cutoff is positioned in the focal point of the OAP and set to reflect 75% of the focused light. The cutoff allows adjusting the contrast of the Schlieren image recorded by a high-speed camera (HS camera) with a resolution of 51.8 μm/pixel and a mounted Nikon zoom lens (70 - 210 mm focal length, f/4).The OAP of Thorlabs have a fixed reflection angle of 15°, a focal length of 381 mm and a diameter of 50,8 mm. They are made out of a silver-coated aluminum substrate and reflect a wavelength range from 450 nm to 20 (cid:586)m. The deflecting mirrors have a flat square reflection surface with an edge length of 50.8 mm. All components are mounted onto breadboards with a surface area of 300 x 300 mm² and firmly connected to the test chamber using aluminum profiles.A VEO640L HS camera from Vision Research generates the high-speed videos. The monochrome sensor enables a maximum image acquisition frequency of 1,600 Hz at full resolution of 2,560 x 1,600 pixels.

The monochrome sensor enables a maximum image acquisition frequency of 1,600 Hz at full resolution of 2,560 x 1,600 pixels. The Schlieren images are captured at a resolution of 1024 x 500 pixels, an image acquisition frequency of 9,656 Hz and an exposure time of 14.64 μs. For the Shadowgraph shots a resolution of 896 x 500 pixels, an image acquisition frequency of 10,568 Hz and an exposure time of 5.13 μs are set. The laser modulation signal is used as a trigger signal.The laser radiation with a wavelength of 1070 nm and a nominal laser power of 400 W is generated by an ytterbium single-mode fiber laser beam source from IPG Photonics GmbH. A HurryScan 20 scanner system with an RTC-5 control card enables the deflection of the laser beam. An F-Theta lens with a focal length of 254 mm focuses the laser beam to a beam diameter dB of approx. 80 μm in the working plane.

An F-Theta lens with a focal length of 254 mm focuses the laser beam to a beam diameter dB of approx. 80 μm in the working plane. This corresponds to the surface of the powder bed.Gas atomized stainless steel powder 316L with a particle size distribution of 22 - 63 μm and an average particle diameter of x50 = 36.86 μm represents the raw material of the process. The powder is applied with a layer thickness DL = 60 μm by means of a silicone lip. The round substrate plates with a diameter of 100 mm also consist of 316L and the corresponding surfaces are roughened with sandpaper and then cleaned with ethanol. Additionally, a 5 mm thick detachable metal sheet manufactured out of 316L is used on top of the substrate plate to facilitate the generation of cross sections.

Additionally, a 5 mm thick detachable metal sheet manufactured out of 316L is used on top of the substrate plate to facilitate the generation of cross sections. The experiments were carried out directly on the substrate plate without a pre-melted, shrunk layer.The noble gas argon is used to create a shielding gas atmosphere by flooding the process chamber until the measured residual oxygen value drops below 200 ppm. By manually controlling the argon supply, the residual oxygen value can be kept below 200 ppm during the L-PBF process. The investigations are performed at an overpressure of 80 mbar relative to the ambient pressure. The changing weather-dependent ambient pressure is not measured. The flow velocity across the substrate plate is set to approx. 2 m/s.

The changing weather-dependent ambient pressure is not measured. The flow velocity across the substrate plate is set to approx. 2 m/s. This value corresponds to the maximum flow velocity in the test setup considered here, at which the metal powder is not removed by the shielding gas flow.To recreate process conditions comparable to industrial parts two scanning patterns are defined. These are rectangular shaped with edge lengths of 35 mm x 6 mm and are schematically depicted in Figure 1. The side view scanning pattern contains 8 stripes with fill vectors aligned1525parallel to the long edge of the rectangle. The six central stripes are 5 mm wide while the two border stripes are 2.5 mm. The tip view scanning pattern contains two 3 mm wide stripes with fill vectors aligned perpendicular to the long edge of the rectangle. The stripes of the side view pattern are exposed from bottom to top and the tip view pattern from left to right.

The stripes of the side view pattern are exposed from bottom to top and the tip view pattern from left to right. Both patterns begin with the start vector of the first stripe and finish with the end vector of the last stripe.Flow direction Flow direction5 mm35 mmSide View PatternStart vectorStripeEnd vectorTip View PatternStripeFill vectorFigure 1. Schematic depiction of the scanning patterns used. Top: Side View Pattern with 5 mm and 2.5 mm wide stripes. Bottom: Tip View Pattern with 3 mm wide stripes.The scanning patterns were chosen to take into account the point of view of the HS camera. The long edge of the rectangle is oriented parallel to the observation plane. Hence, only the side of the vectors are visible, when recording the side view pattern. The tip view pattern enables capturing the tip and tail of the vectors respectively.The exposure parameters used for the experiments were developed by Pichler et al. [6].

The tip view pattern enables capturing the tip and tail of the vectors respectively.The exposure parameters used for the experiments were developed by Pichler et al. [6]. A relative part density of 99.92% is achieved with using PL = 250 W, vscan = 800 mm/s and a hatch distance of Δyh = 100 μm. Hence, these process parameters are utilized for all described experiments in the following.The exposure of the fill vectors is performed in skywriting mode 3 with NPrev = 200 μs,NPost = 200 μs, Timelag = 280 μs and Laser-On-Shift = 20 μs.The ImageJ distribution FIJI allows to preprocess the video recordings [7]. This software also enables to correct darkening of single frames by means of flat field correction which is applied to all images presented in this work. A flat-field-correction is able to correct edge shadowing of the lens, scratches, dust etc. by means of a reference image of a uniformly bright image surface. Each pixel of every single video frame is adjusted in its intensity.

by means of a reference image of a uniformly bright image surface. Each pixel of every single video frame is adjusted in its intensity. Additionally, a tonal value correction is performed with manually determined parameters to achieve the highest possible contrast in the image without losing information. The parameters used are determined separately for Shadowgraph and Schlieren shots, as the different exposure settings lead to different brightness distributions.To analyze the surface and cross sections of the samples contactless measuring devicesare used. Surface imagery is performed through a scanning electron microscope LEO 1455EP1526m m 6m m 3with a magnification of up to 4000x. Colored imagery and analysis of the cross sections are enabled by means of a reflected light microscope from Olympus Deutschland GmbH.Results and DiscussionThis work covers four exposures for result discussions.

Two side view and two tip viewexposures are investigated with and without shielding gas flow respectively.The top images in Figure 2 show exposures of two side view patterns with (a) and without (b) shielding gas flow support. In both cases the melt pool is visibly glowing and shining particles get ejected from the melt pool. The exposure without shielding gas flow support shows an additional glowing area above the melt pool. This is caused by laser irradiation interacting with the particles ejected from the melt pool.With shielding gas flow With shielding gas flowWithout shielding gas flow Without shielding gas flowFlow direction(a)Glowing condensateGlowing particlesMelt pool10 mmFigure 2. In-situ images of the exposure of stripes with (left) and without (right) shielding gas flow support. The process without shielding gas flow is forming a glowing pillar out of condensate and vapor.

In-situ images of the exposure of stripes with (left) and without (right) shielding gas flow support. The process without shielding gas flow is forming a glowing pillar out of condensate and vapor. PL = 250 W, vscan = 800 mm/sThe images in Figure 3 show the resulting surfaces of two corresponding exposures of the tip view patterns. The exposure in the upper image is supported by the shielding gas flow and the lower is not. The surface of the exposure with shielding gas flow shows an evenly distributed pattern of melt tracks. In contrast, the surface without shielding gas flow appears to have an evenly distributed pattern of melt tracks up to 5/6th of the lower stripe until a transition zone. The last sixth of the lower stripe shows this light colored transition zone of approx. 2 mm before changing to a darker color. This darkening and undefined surface structure are also visible on the upper stripe until the end of the exposure on the upper right side.

2 mm before changing to a darker color. This darkening and undefined surface structure are also visible on the upper stripe until the end of the exposure on the upper right side. Light Balling can be observed on the surface after the transition zone that is increasing on the upper stripe.1527(b)Evenly distributed melt track pattern Evenly distributed melt track patternFlow direction2 mmDark surface and balling Dark surface and ballingNo flowExposure directionEvenly distributed melt track pattern Evenly distributed melt track patternTransition zone Transition zoneFigure 3. Reflected light microscope images of two consecutive stripe exposures of tip view patterns with (top) and without (bottom) shielding gas flow. The surface without shielding gas flow shows a transition from a stable process to an unstable darkened process with balling at around 5/6th of the lower stripe in exposure direction. 1: Upper stripe. 2: Lower stripe.

1: Upper stripe. 2: Lower stripe. PL = 250 W, vscan = 800 mm/sScanning electron microscope images are taken and shown in Figure 4 for understanding the process of surface structure formation. Image areas (a) and (b) show consecutive pictures of the lower stripe of a tip view pattern exposure without shielding gas flow. Initially, evenly distributed single melt tracks can be observed until the melt tracks enter a transition zone and become undistinguishable to each other. The evenly distributed melt tracks are similar to the ones in image area (c) that depicts the melt tracks of a section of the top stripe with shielding gas flow. Following the transition zone light balling becomes apparent (b) with a sphere diameter of up to 150 μm. Furthermore, no single melt tracks as in (a) can be identified. The balling effect continues at the beginning of the upper stripe (not shown) and becomes stronger until the end of this stripe (d). The spheres now reach a diameter of up to 250 μm.

The balling effect continues at the beginning of the upper stripe (not shown) and becomes stronger until the end of this stripe (d). The spheres now reach a diameter of up to 250 μm. Powder particles are attached to all surfaces regardless of their overall topology.15281212(a)(b)Single melt trackTransition zoneLight ballingEnd of lower stripe(c)(d)Foreign bodiesPowder particlesStrong ballingc, dabFigure 4. Surface of post-exposure tip view patterns. (a) and (b) show the end of the lower stripe exposure without shielding gas flow with the transition zone and the beginning of light balling. The bottom images show one section of the upper stripe with shielding gas flow (c) and without shielding gas flow (d). In (d) strong balling can be observed. PL = 250 W, vscan = 800 mm/sThe darkening and balling only occur in absence of the shielding gas flow. Residual oxygen in the process chamber was carefully monitored and kept below 200 ppm.

PL = 250 W, vscan = 800 mm/sThe darkening and balling only occur in absence of the shielding gas flow. Residual oxygen in the process chamber was carefully monitored and kept below 200 ppm. Therefore, oxidation is not considered as the cause for the darkening and the increased balling.It was intended in the following to visualize the convection processes of heated gas and the particle movement for finding an explanation on how the absence of shielding gas flow leads to a transition from a stable to an unstable process. In Figure 5 a comparison between Shadowgraph (top) and Schlieren (bottom) images is shown. The Shadowgraph images allow the illustration of particles that block light such as metal condensate and powder particles. The Schlieren images enable the visualization of the convection flow of the heated shielding gas. It needs to be noted that Shadowgraph and Schlieren Images could not be taken at the same time.

The Schlieren images enable the visualization of the convection flow of the heated shielding gas. It needs to be noted that Shadowgraph and Schlieren Images could not be taken at the same time. Hence, the images show the same exposure of a tip view pattern but are recorded in two separate experiments with differing timestamps. The images show the process without shielding gas flow in rows (a) and (c) and with1529shielding gas flow in rows (b) and (d) at three different points in time. A dark plume with an upwards direction can be seen at 283.985 ms into the exposure in the Shadowgraph image (a). The corresponding Schlieren illustration at 283.968 ms in row (c) shows the convection of hot gas with the same upwards direction. The images at 1324.915 ms (Shadowgraph) and 813.328 ms (Schlieren) show the process entering the transition zone described in Figure 3.

The images at 1324.915 ms (Shadowgraph) and 813.328 ms (Schlieren) show the process entering the transition zone described in Figure 3. The plume in (a) becomes lighter in color and the corresponding Schlieren image in (d) depicts a weakening convection flow which indicates a change in process stability. The amount of ejected particles seems to decline compared to the images at 238.985 ms. At 1419.545 ms (Shadowgraph) and 906.914 ms (Schlieren) the transition is completed and light balling begins to occur.

Tip view pattern exposure at start, transition and after transition with light balling. Top: Shadowgraph Images without (row (a)) and with shielding gas flow (row (b)). Bottom: Schlieren without (row (c)) and with shielding gas flow (row (d)). Schematic depiction of scanning pattern at the bottom shows the current exposure position. PL = 250 W, vscan = 800 mm/s1530The corresponding images of the shielding gas flow supported exposures in rows (b) and (d) show no changes in plume behavior. The shielding gas flow pushes the heated shielding gas and entrained particles from the right to the left side and away from the laser beam. This leads to the assumption that the accumulation of particles in the process without shielding gas flow influence the laser beam. The consequences might be the altering of the intensity distribution of the laser beam through scattering as well as the absorption of portions of the light described by Š(cid:171)eglov in [8].

The consequences might be the altering of the intensity distribution of the laser beam through scattering as well as the absorption of portions of the light described by Š(cid:171)eglov in [8]. This assumption is supported by the change of the particle ejection direction during the exposure of the side view pattern shown in Figure 6.(a)Convection front Convection frontCondensate / vaporEjection direction(c)Ejection directionNo flow No flowPoint of view ! Point of viewCurrent exposure position ;; ;; Current exposure positionFigure 6. Side view pattern exposure. (a) and (b): At start. (c) and (d): after transition. Without shielding gas flow. Left: Schlieren Images. Right: Shadowgraph Images. PL = 250 W, vscan = 800 mm/sHoppe et al. [4] analyzed the ejection direction of particles by measuring the ejection angle from the melt pool via particle tracking.

Right: Shadowgraph Images. PL = 250 W, vscan = 800 mm/sHoppe et al. [4] analyzed the ejection direction of particles by measuring the ejection angle from the melt pool via particle tracking. They investigated that at PL = 250 W and vscan = 800 mm/s the particles are ejected against the exposure direction by an average angle of 30° with regard to the powder bed surface. The arrows in (a) and (b) indicate the ejection direction without shielding gas flow at the start of the exposure of the side view pattern. The particles get ejected as described by Hoppe et al. against the exposure direction with alternating orientation. This can be explained by the bidirectional vectors of the scanning pattern described in Figure 1. After passing a transition zone as described in Figure 3 and Figure 4 the ejection direction changes from 30° to almost perpendicular to the powder bed as shown in Figure 6 (c) and (d). This leads to particles being ejected almost directly into the laser beam.

This leads to particles being ejected almost directly into the laser beam. An increase of the ejection angle was also observed by Hoppe et al. [4] when reducing the laser power PL while keeping the scanning speed constant. This supports the assumption that the particles might attenuate the laser beam through scattering and1531(b) (b)(d)5mm 5 mmabsorption of light. The attenuation effect is reinforced by the particles being ejected almost directly into the laser beam. This also explains the observation of increased balling on the surface of the exposed areas in Figure 4. Powder particles are also lifted into the laser beam and subsequently melted midair as depicted in Figure 7.Position of laser beam .... ~ .... • - Position of laser beamFusing particles Fusing particles1 I I1--- 5-q me~mmFigure 7. Time sequence of the fusion of melted particles midair. Side view pattern without shielding gas flow.

Time sequence of the fusion of melted particles midair. Side view pattern without shielding gas flow. PL = 250 W, vscan = 800 mm/sThese melted particles drop onto the solidified surface of the exposed areas forming solid spheres. Some of the melted particles fuse midair and form larger particles with a diameter of up to 250 μm as shown in Figure 4. Part defects are more likely to happen in these areas since the particles are remelted in consecutive layers as described by Tolochko et al. in [9] and Li, Liu et al. in [10].Through Schlieren and Shadowgraph imagery it is possible to visualize the changes in the melting process in dependency of the shielding gas flow. The occurrence of balling is visualized and linked to the change in ejection angle of the particles while the laser beam is altered before reaching the powder bed. The darkening of the surface could not be explained by the Schlieren and Shadowgraph imagery.

The darkening of the surface could not be explained by the Schlieren and Shadowgraph imagery. Since oxidation can be ruled out as a reason it is thought that condensate nanoparticles adhere to the solidified part surface. Particles with a diameter of 80 - 150 nm in the welding plume are described by Š(cid:171)eglov in [8]. These nanoparticles would be small enough to cause the darker appearance of the surface by plasmonic surface resonance. To verify this assumption, additional SEM images were taken with a magnification of 4000x and are shown in Figure 8. Image (a) depicts the surface of a shielding gas supported exposure and (b) the surface of an unsupported exposure.1532-•a)With shielding gas flowb)No flowCondensate nanoparticlesEHT = 20.38 kV13 mm Stage at Y = 41.675 mmEHT =20.38kV13 mm Stage at Y = 34.547 mmFigure 8. SEM images at 4000x magnification. (a) With shielding gas flow. (b) Without shielding gas flow.

SEM images at 4000x magnification. (a) With shielding gas flow. (b) Without shielding gas flow. PL = 250 W, vScan = 800 mm/sThe surface in image (a) shows no sign of adhesion of particles smaller than 1000 nm. Without shielding gas flow on the other hand, the surface is covered with particles smaller than 200 nm, which coincide with the expected size of 80 -150 nm. The sections of the exposed stripes with evenly distributed melt tracks do not show a darkened surface nor nanoparticles. Hence, the presence of the nanoparticles can only be linked to the interaction with the laser beam when the particles are ejected perpendicular to the powder surface. A similar observation has been made by Hoppe et al in [4] showing a powder particle interacting with the laser beam and then being accelerated in direction of the powder bed surface. The reason is the almost instantaneous evaporation of material on the laser beam facing side of the particle leading to recoil pressure.

The reason is the almost instantaneous evaporation of material on the laser beam facing side of the particle leading to recoil pressure. The adhesion cannot be defined as strong since the nanoparticles can be brushed from the surface by mechanical interaction.It was shown that the absence of shielding gas flow leads to an accumulation of nanoparticles in the process atmosphere leading to altered process conditions. This can be observed via Schlieren and Shadowgraph imagery as well as surface topology analysis. To evaluate the influence on the melting tracks we analyzed the shape of the exposure patterns. We expect the melt track width to change after the transition zone since the laser beam characteristics seem to be altered.In Figure 9 detailed surface images of the scanning patterns from Figure 3 are depicted. The boundary of the upper and lower stripe of a tip view pattern without shielding gas flow is shown.

The boundary of the upper and lower stripe of a tip view pattern without shielding gas flow is shown. The melt tracks of the lower stripe are evenly distributed while the upper stripe melt tracks cannot be clearly differentiated from each other. The upper stripe melt tracks are wider compared to the lower stripe melt tracks. The wider appearance suggests a conduction welding process. A similar melt track topology can be seen in (b), which shows the transition zone. The lower stripe melt tracks transition from narrow and evenly distributed melt tracks to wider non-distinguishable melt tracks.1533a)No flowb)No flowUpper stripeUnstable melt trackLower stripeStable melt trackTransition zonec)With shielding gas flowd)With shielding gas flowStable melt track0.5 mma,cb,dFigure 9. Reflected light microscope images of tip view patterns without (a and b) and with (c and d) shielding gas flow.

Reflected light microscope images of tip view patterns without (a and b) and with (c and d) shielding gas flow. (a) and (c) represents the boundary between the upper and lower stripe near the beginning of the scanning pattern. (b) shows the transition zone and (d) the corresponding section of a shielding gas supported pattern.The melt track geometry imply that the laser beam is altered at the transition zone in (b). The wider melt tracks can be explained by a reduction in laser power as well as a change in intensity distribution. Further analysis of the melt tracks via cross sections need to be conducted to verify the assumption of welding mode transition through attenuated laser power.ConclusionThe presented work shows the importance of a reliable shielding gas flow for a stable L- PBF process. Schlieren and Shadowgraph imagery enabled visualizing the convection flows that entrain the nanoparticles emerging from the melting process.

Schlieren and Shadowgraph imagery enabled visualizing the convection flows that entrain the nanoparticles emerging from the melting process. The nanoparticles are ejected into the process chamber depending on the melt pool geometry and are then distributed through natural convection. The absence of shielding gas flow leads to an abundance of nanoparticles in the shielding gas atmosphere in the process chamber. These nanoparticles begin to interact with the laser beam until reaching a tilting point. Entering a transition zone, the melt track topology changes due to the altering of the laser beam possibly by means of absorption and scattering. In the transition from a stable to an unstable process, ejection behavior changes and particles are directly lifted into the laser beam leading to further attenuation and balling on the solidified part surface. A darkened surface through adhered nanoparticles was observed after the transition zone indicating an unstable process.

A darkened surface through adhered nanoparticles was observed after the transition zone indicating an unstable process. The analysis of the melt track topology reinforces the observation made with Schlieren and Shadowgraph images. The melt tracks after the transition zone become wider compared to the melt tracks seen in a stable process with shielding gas flow. Since residual oxygen was kept below1534200 ppm, the most reasonably explanation for this behavior is the reduction of the effective laser power reaching the powder bed and a change of the beam profile.Following this work, further investigations on the altering of the beam profile and effective laser power reaching the powder bed need to be conducted. The concentration of nanoparticles leading to an unstable process are of special interest in a multi laser setup.

The concentration of nanoparticles leading to an unstable process are of special interest in a multi laser setup. Cross sectional analysis of the melt tracks would prove the transition into a conduction welding regime and explain the changes in particle ejection behavior.Funding InformationThis work was supported by the "Research Campus - Public-Private Partnership for Innovation" funding initiative of the Federal Ministry of Education and Research (BMBF) of Germany.References[1] B. Ferrar, L. Mullen, E. Jones, R. Stamp, C.J. Sutcliffe, Gas flow effects on selective lasermelting (SLM) manufacturing performance, Journal of Materials Processing Technology 212 (2012) 355–364.[2] P. Bidare, I. Bitharas, R.M. Ward, M.M. Attallah, A.J. Moore, Fluid and particle dynamics inlaser powder bed fusion, Acta Materialia 142 (2018) 107–120.[3] P. Bidare, I. Bitharas, R.M. Ward, M.M. Attallah, A.J.

 Analysis Technique for Layered Manufacturing Based on Quasi-Wavelet TransformsCheol H. Lee and Charles L. Thomas Department of Mechanical Engineering University of Utah Salt Lake City, Utah 84112SUMMARYAn analysis technique based on the Wavelet transform (WT) has been recently introduced that allows the spatial frequency content of objects produced by layered manufacturing (LM) techniques to be interpreted in terms of manufacturable features. (Lee and Thomas, 1997) Using Haar's wavelet as a basis function, layers with vertical edges are modeled exactly. Using analysis, a 3D model can be transformed, filtered, and inverse transformed resulting in an image of the part as it would look if constructed from layers of a specific thickness. In order to extend this analysis to construction techniques using higher order edges (ruled surface edges or curved edges), the quasi-wavelet transform (QWT) is introduced.

In order to extend this analysis to construction techniques using higher order edges (ruled surface edges or curved edges), the quasi-wavelet transform (QWT) is introduced. QWT analysis is conceptually the same as WT analysis, except that the basis function can be selected by the user, allowing exact analysis of layered manufacturing techniques using higher order construction algorithms. This work is supported by a grant from the University of Utah Research Foundation.INTRODUCTIONWhen a solid model of an object is reproduced physically using a layered manufacturing technique, the result is a double approximation of the electronic model. The first level of This triangular approximation occurs when the model tessellation of the part surface clearly results in error between the CAD model and the triangular mesh. The second level of approximation occurs when the part is constructed from layers.

The second level of approximation occurs when the part is constructed from layers. Using finite thickness layers, any surface that is not either parallel or perpendicular to the slicing plane has a ridged, stepped appearance. Researchers are attempting to address the stepped edge problem by cutting ruled surface edges on the layers (Lee et al, 1995), but the result is still an approximation to the designed surface.is converted to the *.stl format.In order to minimize the error caused by these approximations and maximize the construction speed by increasing the layer thickness, it would be useful to have analysis tools that would assess the manufacturability of a given model using LM. Desirable features of these techniques might include: The ability to compare the geometric complexity of a specific part to a specific devices' ability to create complex geometry. The ability to predict the output of a specific device.

The ability to predict the output of a specific device. The ability to identify the preferred slicing direction for a specific part.A previous paper addresses these issues and suggests that objects be decomposed in terms of spatial frequency. (Lee, 1997) Fourier analysis was shown to be of little use since it does not identify the location on the part of the high or low frequency information. Wavelet analysis was shown to be quite useful since the frequency information is spatially localized and Haar's wavelet provides an exact match to the stepped edge layers produced by existing LM devices.577This paper provides a brief discussion of Fourier and Wavelet Transform Analysis (FTA and WTA) applied to 3D objects and introduces Quasi-Wavelet Transform Analysis (QWA).Fourier AnalysisThe Fourier transform is currently used in applications such as signal analysis and image processing.

It allows frequency analysis and filtering, of an analog signal by decomposing a time domain signal in terms of it's frequency content. A wave form can be filtered and inverse transformed to predict the form of the time domain signal if certain frequencies were missing. Using a 2D FT, images can be analyzed or compressed by deleting the information for frequencies that make little contribution to the image features. (Proakis, 1992)AID FT is often used to examine a time domain signal (amplitude as a function of time). We can examine a slice of an object in terms of spatial frequency by replacing amplitude and time with the two coordinate axes of the slice. Since LM devices in general can create quite complex geometry in the plane of the slice, but are limited by the layer thickness in the slicing direction, one might say these devices are low pass filters in the slicing direction and that they pass all frequencies in the slicing plane.

If basis functions can be selected that represent the cutter shape, then Fourier analysis will produce meaningful results for LM.The limitation of the Fourier transform that lead to the investigation of wavelettransforms is the inability to spatially localize the frequency information. Fourier analysis averages the frequency content over the entire shape under analysis. The transform cannot identify that certain parts ofthe object are simple and other parts are complex.Wavelet AnalysisThe Wavelet transform decomposes a signal in terms of both frequency and position. While it has similar applications to the FT, it not only identifies the frequency content of a signal, but also keeps track of where in the signal the frequencies occurred. The WT decomposes a signal using basis functions called wavelets.

The WT decomposes a signal using basis functions called wavelets. A wavelet is a family of functions derived from a single function, called mother wavelet expressed as:w,.,,(x)= ~"t:b)where 'a' and 'b' adjust the frequency and position of the wavelet. Using the WT, a signal, f(x) can be expressed as: (Newland, 1993, and Tewfik, 1992)W(4x) W(4x-l) W(4x-2) + ...W(4x -3)and578(I)(2)U(X)={~for O~x <1 otherwiseis a contribution of a wavelet at frequency 2J with a where a j shift k. At a frequency 2J , there are 21 wavelets with independent amplitudes. Fig. 1 shows four levels of wavelet components for a typical transform. The wavelet used here is Haar's D2 wavelet which consists of a step to amplitude ai followed by a step down to ai. Thus the WT to spatially localizes the frequency information.2 +kFig.

The wavelet used here is Haar's D2 wavelet which consists of a step to amplitude ai followed by a step down to ai. Thus the WT to spatially localizes the frequency information.2 +kFig. 1: Wavelet componentsfor a typical transform.Although wavelets must meet criteria similarto theorthogonality constraints required for basis functions for FT, one wavelet, Haar's D2 wavelet function, consists of steps. Thus, it is perfect for representing the stepped edges of LM. Using Haar's D2 wavelet, the levels of the WT correspond to layer thickness'. The summation of several wavelets of different frequencies results in a stepped surface that can always be reproduced using layers corresponding to the highest of the summed frequencies. Using WT it is possible to transform a 3D object, filter the object, and inverse transform the object producing a prediction of the geometry realized from layers of a selected thickness. (See Fig.

Using WT it is possible to transform a 3D object, filter the object, and inverse transform the object producing a prediction of the geometry realized from layers of a selected thickness. (See Fig. 2) (Lee, 1997) Note that the object is frrst mapped into cylindrical coordinates to get a single valued function.There are limitations to the wavelet transform: The wavelets double in frequency at each level. Thus, only certain layer thickness' are actually considered in the analysis. If a LM device were capable of varying the layer thickness in an analog fashion, the technique as described would not be able to properly analyze the problem. A second limitation is that there are no known wavelets that match the cutter shapes used by the new higher order construction algorithm LM devices that are currently under development.QUASI-WAVELET ANALYSISThe Quasi-Wavelet Transform follows the spirit of the WT, but relaxes the constraints on the basis functions.

As a result the QWT can use any function as a basis function. The QWT decomposes a function in terms of quasi-wavelets that are scaled and shifted in a similar fashion to wavelets. The requirement for QWT is that, at the lowest frequency, a single quasi-wavelet is scaled such that the squared error between the quasi-wavelet and the function is minimized. The next level has two quasi-wavelets and each is scaled independently to minimize the squared error between the function and the sum of the two quasi-wavelet levels. At the third level there are four wavelets and each is scaled to minimize the error between the function and the sum of the three levels of wavelets. This process continues until the highest frequency of interest is reached.A function W(x), made up of a linear combination of general functions g(x), is defmed as a quasi-wavelet function, within the interval O<=x<=1, and is set to zero outside the interval.

Complex objects with multiple surfaces must be analyzed one580(4)(5)(6)(7)~J-l~l~-lJ;;-lx 00 Fig. 3: This shows three quasi-wavelet basis functions.x 00x ~a) Step - This is equivalent to Haar's 02 wavelet. b) Angle - This quasi-wavelet models the ruled cut algorithm for layered manufacturing. c) Curve This quasi-wavelet models the curved cut algorithm for layered manufacturing.surface at a time. The z axis of the cylindrical coordinate system is assumed perpendicular to the slicing plane. When applying QWT, the quasi-wavelet function should be carefully chosen so the basis function can properly represent the LM machine. Quasi-wavelet functions for stepped edges (Stepped Cut Algorithm, SCA), ruled edges (Ruled Cut Algorithm, RCA), and curved edges (Curved Cut Algorithm, CCA), were selected as follows:WSCA(z)=Jf;(z)=al u(z)WRCA(z)=~(z)=alu(z)+a2z WCCA(z)=~(z)=al u(z)+a2z+a3Z2where aj are the coefficients that define the shape of each quasi-wavelet.

An example of each is shown in Fig. 3. The QWT coefficients, aj's, of Wn(z) for SCA, RCA, and CCA for all levels, make up {ai} as follows:where the first subscript identifies the level (frequency), the second subscript identifies the position of the wavelet, and the third subscript identifies the function within the wavelet. Fig. 4 is a pictorial example of the coefficients for QWTRcA at level I and 2. In QWTSCA' if there are 2m 1 + coefficients, then all is the average value for all data, and aij is the coefficient of u(2 i z - j + I). 2 coefficients in the QWTRCA, am is the data average, and a1l2 is usually zero, If there are 2m + where aijl is the coefficient of u(2 i z- j+l) and aij2 is the coefficient of (2 i z- j+I).

2 coefficients in the QWTRCA, am is the data average, and a1l2 is usually zero, If there are 2m + where aijl is the coefficient of u(2 i z- j+l) and aij2 is the coefficient of (2 i z- j+I). In a QWTCCA, if there are 3 x 211l+1 coefficients, am is the data average, and a1l2 and am are usually0.30.30.1 a111,a112a12I,a 1220.1 211,a212::E::E0.10.1a241,a2420.30500 n (a)10000.30500 n(b)1000Figure 4:This shows two levels of a transform using the Angle quasi-wavelet. each sloped quasi-wavelet has its own independent coefficients.581(8a)(8b)(8c)zero, where aijI is the coefficient of u(2 i z- j+l), aij2 is the coefficient of(2 i z- j+l), and aij3 is the coefficient of (2i z - j +Ii. APPLICATIONSFrequency AnalysisIn LM, expressing objects in the frequency domain is very meaningful because frequency can be related to the 0.0 number of layers.

APPLICATIONSFrequency AnalysisIn LM, expressing objects in the frequency domain is very meaningful because frequency can be related to the 0.0 number of layers. The contributions of individual frequencies to the geometry of a part can be found using frequency analysis. At certain cut off frequencies, the loss of geometric information is easily visualized through frequency component charts.:-----------+--+-0.40.20.23 4 5Fig. 5 shows the QWTRcA. of a cross section of a vase shown in Fig. 6. The vase has been scaled to unit height Figure 5: This is the Quasi-Wavelet Transfonn of a slice of a with height shown as the z axis. The level axis corresponds to frequency or number of layers with 2° layers at level 'n'. The magnitude is simply the quasi-wavelet components drawn on the M, z axis in this figure. The figure shows higher frequency components near the bottom and top of the vase.

The magnitude is simply the quasi-wavelet components drawn on the M, z axis in this figure. The figure shows higher frequency components near the bottom and top of the vase. The center of the vase has higher amplitude components at low frequency and little contribution from higher frequencies. This suggests that the base can be constructed from 128 layers corresponding to level 7.1 0 1 26 7 8 9levelvase using Angle quasi-wavelets.Output Prediction The output of a SFM device can be predicted by applying a low pass filter to the data from Fig. 5. This is done by setting the magnitudes of all wavelets at levels above the filter cut off to zero. Next, an inverse transform is performed and an image of the object is produced as it would look if constructed using the selected algorithm at the selected layer thickness. Fig. 6 shows examples of output prediction for SCA, RCA, and CCA.

Fig. 6 shows examples of output prediction for SCA, RCA, and CCA. As would be expected, the RCA and CCA produce an accurate representation using thicker layers than required for SCA.Advanced Filtering Using QWT it is not only possible to filter frequencies, it is possible to create filters that match machine constraints. If an RCA device can cut angled edges on the layers, but the angle is limited, this can be accounted for in the transform. A CCA device might be able to produce curved surfaces as long as the radius of curvature is greater than some minimum.For Shapemaker II, a LM machine that implements RCA, hardware limitations constrain the maximum angle of cut. The design of Shapemaker II allows for a maximum cut angle of 45 In degrees. For some layer thickness', the maximum angle may be reduced from this value. QWTRCA, cut angles are specified by the coefficient of the z term. By setting a limiting value on the z terms, it is possible to limit the angle of cut.

QWTRCA, cut angles are specified by the coefficient of the z term. By setting a limiting value on the z terms, it is possible to limit the angle of cut. (Lee, 1997)582M"0.30.20.103~0.2 0.1 QOn0O.3t :g 0.3E::SJ 03r:=:g O.3r:=:gQ2 0.1 0.0Q2 0.1 0 0.0----"S'*'() nn0.3~0.2 0.1 0 . 000.20.1 0 .00 nnO.3CQ0.2 0.1 0 . 00 n0.2 0.1 0 .00 nFigure 6:The vase shape has been transformed using Step, Angle, and Curve quasi-wavelets. Then it was filtered at different frequencies and inverse transformed.DemonstrationThe head shown in Fig. 2 was constructed using Shape Maker II. The model, in STL fonnat was sliced in 1.8 mm thick slices. The slices were cut from commercially available 1.22 m by 2.44 m, white polystyrene foam, with an average thickness of approximately 1.8 cm. The size of the 47 layer head is 57 cm by 74 cm by 83 cm.Using QWTRCA, the shape of the head was predicted for 47 layers with a maximum cutter angle constraint of 45 degrees. Fig.

The size of the 47 layer head is 57 cm by 74 cm by 83 cm.Using QWTRCA, the shape of the head was predicted for 47 layers with a maximum cutter angle constraint of 45 degrees. Fig. 7(a) shows the inverse QWTRCA of a three dimensional head with 47 layers and a 45 degree constraint on the cut angle. Fig. 7(b) shows a completed head made by SM2. As demonstrated by the predicted head, because of machine constraints, there is obvious error due to the constrained angle near the top of the completed head.Advantages and LimitationsThe QWT uses any linear combination of functions as a basis function. Frequency localization identifies where on an object the frequency components exist. As a result, QWT should be capable of analyzing any LM technique. Using the QWT, the output of LM systems can be predicted and reconstruction errors can be evaluated mathematically.

As a result, QWT should be capable of analyzing any LM technique. Using the QWT, the output of LM systems can be predicted and reconstruction errors can be evaluated mathematically. Since the quasi wavelet transfonn decomposes a three dimensional surface geometry into frequency components, a frequency analysis can be perfonned by inspecting the transfonn coefficients. Optimal construction directions can be found through coordinate rotation and frequency analysis.Limitations for the application of three dimensional object analysis mainly arise from limitations of the three dimensional the cylindrical representation method is used, which maps the three dimensional surface of an object onto a cylinder. This is done to produce a single valued function for the analysis. However, this method may still produce a multiple valued function for certain types of geometry's.surface representation. Currently,By nature of the analysis, frequencies analyzed by the QWT are integer powers of two.

Currently,By nature of the analysis, frequencies analyzed by the QWT are integer powers of two. As a result, building an object from 3 layers or 115 layers is not considered. Most current SFM devices build parts from layers of fixed thickness and, as a result, seldom build parts from layers5830Figure 7: The QWT prediction ofthe Shapemaker II process is shown compared to a model constructed usingShapemaker II.that are an integer power of two. This problem can be handled by zero padding. Essentially the model is placed in build volume that is an integer power of two layers thick at the desired layer thickness, and the analysis is done on the entire volume.CONCLUSIONSThe quasi-wavelet transform is a method for mathematically analyzing the spatial frequency content of three dimensional objects. This technique allows the complexity and manufacturability of a part to be analyzed.

This technique allows the complexity and manufacturability of a part to be analyzed. Using an inverse transform it is possible to reconstruct the part as it would look if it were manufactured by a specific process. The quasi wavelet transform is a generalization of the wavelet transform technique that allows any function to be used as a wavelet. This allows any manufacturing process to be modeled by this technique.REFERENCESLee, C. H., Gaffney, T. M., and Thomas, C. L. (1995) Paradigms for Rapid Prototyping, Proceedings of the Sixth International Conference on Rapid Prototyping, pp. 131-143.Lee, C. H., Gaffney, T. M., and Thomas, C. L. (1996) Soft Tooling for Low Production Manufacturing of Large Structures, Proceedings ofthe 1996 Solid Freeform Fabrication Symposium, pp. 207-214.Lee, C. H. and Thomas, C. L. (1997) Wavelet Transfonn Based Analysis of Objects for Layered Manufacturing, Proceedings ofthe Seventh International Conference on Rapid Prototyping,.Lee, C. H.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAnalytical and experimental characterization of anisotropic mechanical behaviour of infill building strategies for fused deposition modelling objectsProf. Marlon Wesley Machado Cunico, Ph.D. , marloncunico@yahoo.com.br Universidade de São Paulo campus São Carlos Concep3D R&DEng. Jennifer Desiree Medeiros Cavalheiro University Federal of ParanaAbstract: As consequence of the 3d printing extraordinary rising along the last years, product development fields are facing new challenges. In addition, it is notory that low cost additive manufacturing, as such fused filament fabrication (FFF), result in objects with anisotropic mechanical behaviour. Nevertheless, there is still lacking studies that highlight a proper specification of those mechanical proprieties.

Nevertheless, there is still lacking studies that highlight a proper specification of those mechanical proprieties. For that reason, the main goal of this work is to present a mechanical characterization of anisotropic behaviour of FFF objects as a function of infill strategy using a finite element method. In this work, the main effect of building parameters were investigated in addition the identification of generalized elasticity and failure mode formulations. By the end, the general reccomendation for objects building was sketched in order to support new strength based developments.Key words: Additive manufacturing, Finite element Method, Anisotropy1. INTRODUCTIONAlong the last years, additive manufacturing, which is also known as 3D printing technologies, has started playing a new role in several market segments.

INTRODUCTIONAlong the last years, additive manufacturing, which is also known as 3D printing technologies, has started playing a new role in several market segments. This might be a consequence of popularization of low cost technologies, which is majorly characterised by fused filament fabrication (FFF) technologies (CUNICO, 2012; CUNICO, 2015).In spite this extraordinary popularization, several challenges still need to be overcome, whereas several product and fabrication issues has been highlighted. Among these problems, material mechanical anisotropy directly affects mechanical strength, possibly jeopardizing functional parts (CHEN et al., 2013; GUESSASMA et al., 2016).In addition, several studies have been done in order to better understand the anisotropic behaviour of FFF objects and fabrication parameters. For example, AHN et al.

et al., 2010),(SOOD, A. K. et al., 2010) and(TOO et al., 2002).Nevertheless, there are few works related to the correlation between process parameters andmechanical strength of FFF objects in a computational environment.For that reason, the present work exposes the theoretical analysis of mechanical anisotropy of FFF by the application of finite elements analyses. In addition, this work generalised the anisotropy984behaviour of and generates numerical formulation which describes the elastic behaviour of FFF parts as a function of bead orientation, air gap and layer thickness.For that study, we have applied a multivariable full design of experiment (24) with 2 levels and no central point where raster orientation (α), distance between lines (d) and layer thickness (h), bead width (w) were the control factors.

In order to compare this work with previous studies, we have also analysed the mean stress (based on specimen dimensions) at break. The virtual experiment design is presented in Table 1, where the variable levels and their values are shown. Table 1 - Virtual Experiment DesignIn this table, values of distance between layers, layer height and orientation and bead width wereshown and the design of experiment we applied in this study was Full (24) with no centre points.For the finite element analysis, we modelled the specimens and fabrication filament, reproducing the main fabrication characteristics which effects the mechanical anisotropy. We have also used the standard specimen shape of ASTM 638 type IV in order to compare the results with exist data.

We have also used the standard specimen shape of ASTM 638 type IV in order to compare the results with exist data. In Figure 1, the transversal section of two specimens were put as example of the filament modelling.985Figure 1 Top view of specimens with a) 0° of orientation bead and 0.35mm of distance between layers b) 0° of orientation bead and 0.45mm of distance between layersFor this analysis, we have also ignored effects of deposition temperature, environment temperature, bed temperature, warpening, and bead width variations. In this way, we can identify the effects orientation, height and distance between layers on the object elastic behaviour.For the finite elements analysis, we defined a tetrahedral mesh with width of 0.025mm.

An example of solid mesh and interaction between bead and layers in transversal cross section is presented in Figure 1.Figure 2 Mesh parameter and interaction between bead with 0.41 mm and layers height of 0.3mmIn order to evaluate elastic behaviour of these specimens, we analysed the strain in xyz directions and shear strain in xy, xz, yz directions, in addition to plane strain. For the analyses of young modulus and Poisson ratios, load combinations were applied and responses were acquired in notch area in order to generate strain-stress (S-S) diagrams.3. RESULTS AND DISCUSSIONS With respect to obtained results, we firstly identified the internal stresses of specimens and compared with average stresses in S-S diagram in the elastic domain of material. In this case, the average stress is relative to external area of notch cross section, simulating ASTM 638 procedures in Laboratory environment. In Figure 3, the S-S diagram of experiment 1 is presented.

In Figure 3, the S-S diagram of experiment 1 is presented. The maximum internal principal stress and the average stress as a function of strain exposes a divergence that is mainly caused the internal porosity of specimen.986Figure 3 - S-S diagram of maximum internal stress versus average stress of experiment 1 (d- 0.4;h-0.2;α-90°;w-0.5)This figure also points that the yield stress of 40MPa of ABS in internal principal stress leads to 20 MPa in the average stress. Therefore, if we assume that yield stress is the design limit, we can identify that the maximum stress that objects would support is 20 MPa.

Therefore, if we assume that yield stress is the design limit, we can identify that the maximum stress that objects would support is 20 MPa. For that reason, we selected this value as one of the main parameters for the virtual experiment analysis.Another parameter that we extracted from this figure is the young modulus of average stress,which is obtained by the derivate of S-S curve.Therefore, we can identify the main and secondary effects of processes parameter on yield stress and young modulus, as presented in Figure 4. In this figure, it is possible to see that orientation is the factor that most contribute for the increase of yield stress and elasticity modulus.Another point that is also possible to see is that the secondary effects are small, indicating thatthe interaction between factors is low.On the other hand, the Poisson ratio in z direction was shown to be most affected by layer height and orientation, as presented in Figure 5.

This figure also indicates that Poisson ratio in y direction is most affected by bead width and distance between beads.Another point that was also exposed in this analysis is that the interaction between control factors is high for Poisson ratio in y direction, where the distance between bead interacts with orientation.

In addition, applying this formulation in Finite element Software, we can optimize the fabrication parameters in order to improve mechanical properties.It was also importance to note that this method indicated the maximum object stress that isacceptable in design as a function of process parameters.Nevertheless, further studies are still needed to better understand applications, implications andcontribution of this method for product and mechanical design.4. CONCLUSIONS In conclusion, this work evidenced the anisotropic behaviour of FFF specimens in acomputational simulation environment.The main and secondary effects were also exposed, where the bead orientation most affects the average yield stress and elasticity modulus. In contrast, bead width and distance between layers most affects Poisson ratio in y direction.

In contrast, bead width and distance between layers most affects Poisson ratio in y direction. The maximum design stress was also shown as a function of process parameters .By the end, a generalized formulation of elasticity was identified, where bead width, distancebetween beads, layer height, and bead orientation were the variables that control the equation.Moreover, more studies are still needed to better understand the anisotropic behaviour inaddition to consider other process parameters in the analysis.5. Acknowledgement The authors would like to thank the CNPQ for Financial support in addition to USP and Concep3D R&D for the support and infrastructure.6. REFERENCESAGARWALA, M. K., V. R. JAMALABAD, et al. Structural quality of parts processed by fused deposition. Rapid Prototyping Journal, v.2, n.4, p.4-19. 1996.AHN, S., M. MONTERO, et al. Anisotropic material properties of fused deposition modeling ABS. Rapid Prototyping Journal, v.8, n.4, p.248-257.

In this matter, the complex thermal history during processing often states a remaining uncertainty. Besides in situ measurements and numerical estimations, analytical models can be used to achieve a deeper understanding of the transient temperature evolution. In this work, an iterative solution to the analytical Rosenthal equation is developed and applied to ZrCuAlNb- and CuTiZrNi-BMGs to predict melt pool dimensions and cooling rates during PBF-LB/M. Therefore, temperature-dependent thermal properties are determined via laser flash measurements. The effective absorptivity of the two materials is measured, and single-line experiments were performed as a validation for the approach.Keywords: Bulk Metallic Glasses, Additive Manufacturing, PBF-LB/M, Rosenthal ModelIntroductionAdditive manufacturing (AM) is widely recognized for its high design freedom, individualization capabilities, and lightweight design.

Recently possibilities to manufacture novel advanced materials are brought into focus. Among the various AM technologies, laser powder bed fusion (PBF-L/BM) has reached a significant level of technological maturity for metallic parts. The experience with standard materials is high, and applications on an industrial scale are already being implemented [1]. However, the exploration of new materials that can646fully benefit from the specific process characteristics – i.e., layer-wise buildup, high heating and cooling rates – is still subject to research. One promising alloy class for the processing via PBF-LB/M is the bulk metallic glasses (BMGs) [2, 3]. These materials exhibit higher strength, elasticity, and superior cryogenic properties compared to crystalline alloys [4].

These materials exhibit higher strength, elasticity, and superior cryogenic properties compared to crystalline alloys [4]. The emergence of an amorphous structure in a metallic alloy requires rapid cooling with an alloy specific so called critical cooling rate (Rcrit).In conventional processing routes, these rates are typically difficult to achieve with increasing volume to surface ratio, which effectively limits the manufacturable size of BMGs to several millimeters [5]. In contrast, PBF-LB/M inherently features high cooling rates of 10^6 to 10^7 K/s [6], which exceeds the Rcrit of most BMGs by several orders of magnitude. This theoretically enables the manufacturing of large and complex BMGs. Nevertheless, in practical applications, partial crystallization is frequently observed [7–13]. The layer-wise deposition of powder and vector-based heating implies a heat treatment of the surrounding solidified material.

The layer-wise deposition of powder and vector-based heating implies a heat treatment of the surrounding solidified material. Additionally, contaminations of atomized powders and the influence of the process gas are considered to affect the structure of fabricated materials [14, 15]. Consequently, to ascertain vitrification despite the uncertain limitations, alloys with comparatively high glass-forming ability (GFA) are preferably chosen.Despite the potential of PBF-LB/M, the reported high cooling rates suggest that its capabilities have not been fully exploited yet. The mechanism underlying crystallization and vitrification is not yet fully understood, and experimental determination of the complex thermal cycling during additive manufacturing processes on a meso scale cannot easily be performed. Analytical methods for determining cooling rates and melt pool dimensions could enable an estimation of the processability, including vitrification of novel material classes with lower GFA.

Analytical methods for determining cooling rates and melt pool dimensions could enable an estimation of the processability, including vitrification of novel material classes with lower GFA. Further, the choice of adequate process parameters could be based on the analytically determined viable cooling rates and melt pool dimensions, enabling an accelerated investigation of novel and processable BMGs.Analytical Rosenthal modelAn analytical approach to describe the heat dissipation during laser welding was developed by Rosenthal in [16] (eq. 1). It is based on a moving point heat source as energy input, constant material parameters, and conduction as the only heat loss:T = T0 +aPL 2πkδe−vs (δ+ξ) 2τEq.

1). It is based on a moving point heat source as energy input, constant material parameters, and conduction as the only heat loss:T = T0 +aPL 2πkδe−vs (δ+ξ) 2τEq. 1with the coordinates x (in scanning direction), y (width of scan track), and z (depth of the melt pool); T0 = temperature of surroundings, T = temperature at a specific point (x|y|z), a = absorptivity, PL= laser power, k = thermal conductivity, vs = speed of heat source (moves along x), τ = thermal diffusivity, ξ = x − vs ∙ t is the distance of point (x|y|z) to heat source in x direction. δ = √ξ2 + y2 + z2 is the radial distance of point (x|y|z) to the heat source.Despite the simplifications, several studies demonstrated the viability of the Rosenthal model to approximate the temperature distribution during PBF-LB/M [17–19]. For instance, Promoppattum et al. used the Rosenthal equation to predict the dendrite spacing of IN718 in [17].

For instance, Promoppattum et al. used the Rosenthal equation to predict the dendrite spacing of IN718 in [17]. In PBF-LB/M the main heat transfer mechanism changes with increasing energy input from conduction mode (dissipation into surrounding solid material) to keyhole mode (evaporation and plasma) [20]. One can conclude that the validity of the model is higher for647processing parameters in conduction mode. Since crystallization needs to be avoided in processing BMGs, low energy densities are typically preferred [13, 21]. Consequently, keyhole mode is generally avoided in BMG processing [22], and the Rosenthal equation potentially allows meaningful estimations for the apparent temperature profiles. Additionally, latent heat and phase transitions are presumably irrelevant during vitrification, further enhancing the practicality of the approach. However, there are remaining uncertainties regarding the assumption of constant thermophysical properties and the unknown absorptivity.

However, there are remaining uncertainties regarding the assumption of constant thermophysical properties and the unknown absorptivity. Metallic materials typically feature large temperature dependencies of their thermal conductivity and diffusivity. It is reasonable to assume that BMGs also display similar temperature dependencies, which should be taken into consideration when applying the analytical model.This work targets the analytical description of the melt pool dimensions and cooling rates during PBF-LB/M of BMGs based on the Rosenthal model. Thereby a fundamental understanding of the time-temperature regime present during the laser material interaction is established. Temperature dependencies of the thermal conductivity and diffusivity are the Rosenthal equation. Two glass-forming alloys, determined and Zr59.3Cu28.8Al10.4Nb1.5 (AMZ4, trade name: Zr0) and Cu47Ti34Zr11Ni8 (Vit101) are chosen for the analysis.

Two glass-forming alloys, determined and Zr59.3Cu28.8Al10.4Nb1.5 (AMZ4, trade name: Zr0) and Cu47Ti34Zr11Ni8 (Vit101) are chosen for the analysis. AMZ4 is a commercially available glass forming alloy, whereas Vit101 is known and investigated, but not yet commercialized. They inhere outstanding material properties, such as an elasticity exceeding conventional materials by ~2%, combined with a flexural strength of 2.1 GPa and 2.5 GPa. The selection of these materials allows for a comparison between analytically predicted and experimentally determined melt pool widths, aiming to validate the model. The novel glass-forming alloy Ti60Zr15Cu17S8 (Ti60S) is a prospect for PBF-LB/M with an extraordinary yield strength of ~3 GPa, exceeding conventionally manufactured Ti6Al4V by 200%. Additionally, biocompatibility and corrosion resistance are favoring this material.

Additionally, biocompatibility and corrosion resistance are favoring this material. The feasibility study concerning manufacturing amorphous Ti60Zr15Cu17S8 single tracks by PBF-LB/M completes this work.insertedintoMaterials and MethodsBMG AlloysFor the three chosen glass-forming alloys AMZ4, Vit101, and Ti60Zr15Cu17S8,the physical properties are given in Tab. 1. Gas atomized powder feedstock was used for the processing of single lines of AMZ4 and Vit101.As a measure for the GFA, the critical diameter 𝐷𝑐 [cm], the highest possible diameter at which the whole structure solidifies amorphous during casting, is present. Based on this, the critical cooling rate in K/s [23] is:𝑅𝑐𝑟𝑖𝑡 =10 (𝐷𝑐)²Eq. 2This is used as a reference value for the comparison to the calculated cooling rates in additive manufacturing:𝑅 =𝑇𝑙 − 𝑇𝑔 𝑡𝑐Eq.

2This is used as a reference value for the comparison to the calculated cooling rates in additive manufacturing:𝑅 =𝑇𝑙 − 𝑇𝑔 𝑡𝑐Eq. 3648Table 1: Composition and physical properties of selected glass-forming materials.Physical property Composition [at%] Density Liquidus temperature Tl Solidus temperature Ts Glass transition temperature Tg Crystallization temperature Tx Critical casting diameter DC Critical cooling rate 𝑅𝑐𝑟𝑖𝑡 (acc. to eq. 2)Ti60S ( [24]) Zr59.3Cu28.8Al10.4Nb1.5 Cu47Ti34Zr11Ni8 Ti60Zr15Cu17S8AMZ4Vit1016.68 g/cm³ [147] 1193 K [26] 1143 K [27] 671 K [26]6.89 g/cm³ [25] 1168 K [25] 1067 K [25] 690 K [25]5.5 g/cm³ 1365 K - 675 K748 K [26]710 K [25]752 K5-14 mm [28] 5 - 40 K/s4-5 mm [23, 25] 20 - 62 K/s1 mm 1000 K/sSample analysisThe thermal conductivity was measured by the laser flash method.

Measurements were performed with a NETZSCH LFA 457 MicroFlash to assess the temperature-dependent thermal conductivity by measuring the thermal diffusivity [29]. Vit101 samples, in the form of pellets, possess a diameter of approximately 10 mm and a thickness of 3 mm. They were coated with graphite on both surfaces their absorption/emission characteristic. The measurements were carried out in an argon atmosphere (purge flow of 75 ml/min) within a temperature range from room temperature to 1073 K, in temperature steps of 100 K. Additional measurement steps were chosen at 623, 723, and 753 K to cover the phase change. Each sample was measured twice. Stepwise heating with 0.02 K/s and holding periods of ~280 s is applied. The specific heat capacity was determined using an equation proposed by Kubaschewski [30] and fitting parameters by Bochtler [25], assuming a crystalline material behavior. The thermal conductivity was calculated employing the procedure introduced by Parker et al.

The thermal conductivity was calculated employing the procedure introduced by Parker et al. [31] using an improved Cape-Lehman model [32] within the analysis software provided by the LFA manufacturer (NETZSCH LFA Analysis version 7.1.0).to improveFinally, the absorptivity is required for the Rosenthal equation. The energy absorption during the laser material interaction in PBF-LB/M has been extensively discussed in the literature over the last decade. One uncertainty is the coexistence of bulk, powder, liquid, and plasma within the interaction zone. This renders conventional techniques such as measurements with an Ulbrich-sphere challenging during the process. Attempts for ray-tracing simulations [33] and analytical calculations based on Hagen-Rubens [34] typically neglect the impact of fume and gas flow during the PBF-LB/M process. Results from Ye et al. imply that the liquid melt is the main absorbing medium.

Results from Ye et al. imply that the liquid melt is the main absorbing medium. However, obtaining the necessary physical properties of liquid metallic melts for analytical or numerical calculations are difficult to obtain and rarely reported. An alternative approach to determine the energy coupling during PBF-LB/M is the calorimetric determination as initially presented by Trapp and adapted by Schnell et al. in [35] and [36]. This setup is used to determine the effective absorptivity of AMZ4 and Vit101. The649absorptivity of Ti60Zr15Cu17S8 is determined by the application of eq. 4 which is derived from the Rosenthal equation (eq. 1):a ≈2D2∗ρ∙Cp∙v∙(Tl−T0)∗πe 8∗PEq. 4where 2D is the weld track width of single tracks, processed in a PBF-LB/M machine on cast material, due to the lack of atomized powder in this feasibility study.PBF-LB/M processingFor empirical validation, single lines are processed via PBF-LB/M.

For the sake of comparison and transferability of the results, two alloys and different manufacturing systems are applied. Argon atomized powder of the BMG-forming alloy Cu47Ti34Zr11Ni8 (in -at.%), with a particle size of 20-63 µm (here called: Vit101), described in detail in [15], and commercially available Zr0 from Heraeus AMLOY Technology (here called by its former name: AMZ4) are used for this study. PBF-LB/M experiments on Vit101 are carried out using the SLM 280 HL system from SLM Solutions, equipped with a 700 W fiber laser and a nominal spot size of 70 µm. Samples are processed with a layer thickness ds = 20 µm and, if applicable, hatch ℎ of 90 µm. AMZ4 is processed on an M100 eos with a nominal laser spot diameter of 40 µm and a maximal laser power output of 200 W. The experiments are conducted in Argon atmosphere (Arcal Prime, 99.998 vol.%). The individual parameters are listed in Tab. A-2 in the appendix.

The experiments are conducted in Argon atmosphere (Arcal Prime, 99.998 vol.%). The individual parameters are listed in Tab. A-2 in the appendix. The line energy density EL according to Meiners is used as equivalent for the energy input at different parameter combinations according to eq. 5  [37]:EL =P vEq. 5Results and discussionThermophysical propertiesAs an essential input variable for the Rosenthal equation (see eq 1.), the thermal conductivity of the investigated alloys is analyzed initially. Fig. 1 illustrates the thermal diffusivity τ plotted over the temperature from 293 K to 1073 K of the initially amorphous samples (run 1) and the partially crystallized sample (run 2). The thermal conductivity is shown in Fig. 2 respectively. Additionally, the coefficients described by Lindwall [38] are depicted for AMZ4 as a reference.Vit101 = 5.17 W/(mK) AMZ4 = 5.7 W/(mK) until the temperature reaches 674 K.

2 respectively. Additionally, the coefficients described by Lindwall [38] are depicted for AMZ4 as a reference.Vit101 = 5.17 W/(mK) AMZ4 = 5.7 W/(mK) until the temperature reaches 674 K. Here, the discontinuous rise of and k295,4 K the coefficient can be associated with the glass transition and crystallization of the samples. The values roughly match Tg of each respective material. During the second heating procedure, a continuous progress of the thermal conductivity indicates the absence of a phase transition, thus implying a crystallized sample. The distinct difference in the heat conductivity between the amorphous and crystalline phase agrees with the result of Yamasaki et al. and originates from the highly alloyed solid solution [39].Values of the amorphous samples increase steadily starting from k295,4 K650Figure 1: Thermophysical properties: Thermal diffusivity of first and second heating cycle and fit function. Left: AMZ4. Right: Vit101. Acc.

Left: AMZ4. Right: Vit101. Acc. to: [40]In view of the large temperature dependence of the thermal conductivity, an iterative solution of the Rosenthal equation based on temperature-dependent values is motivated. Thus, fit functions are determined for both values. During the measurement by laser flash application, the heating rate caused a crystallization of the samples, This crystallization process differs from the fast heating and cooling rates experienced during PBF-LB/M processing, where such crystallization is not expected to occur. Thus, the progression of the values for the transition of amorphous materials to melt is considered. Based on the measured values for the amorphous materials, the fit function should describe this material behavior. The resulting fit-functions are depicted in Fig. 1 and 2 are given in Tab. 2.Figure 2: Thermal conductivity in dependence on the temperature, resulting from Laser Flash measurements. Left: AMZ4. Right: Vit101. Acc.

1 and 2 are given in Tab. 2.Figure 2: Thermal conductivity in dependence on the temperature, resulting from Laser Flash measurements. Left: AMZ4. Right: Vit101. Acc. to: [40]Finally, the absorption coefficient is required for the Rosenthal equation. The time- temperature proceeding is obtained by thermocouples during the processing of a 10 mm disc in three consecutive layers. Applying the procedure described above, the effective absorptivity can be averaged to 0.32 for AMZ4 and Vit101. However, it is important to note that the651absorptivity may change based on the processing conditions and applied process parameters [36].Table 2: Fit functions and absorptivity for AMZ4 and Vit101 used in the following sections.Th. Diffusivity τTh.

Naturally, the melt pool length and the associated time-Temperature (t-T) curve are elongated assuming the thermophysical properties at room temperature. Under the given conditions, the melt pool length is increased by 19.9 % and the width of the HAZ where Tg is exceeded by 22.4 %.Figure 3: Results of conventional Rosenthal model for AMZ4 with k and 𝜏 taken at 293 K and at 1073 K; at P=60 W and 𝑣𝑠=1.4 m/s. Left: T-t diagram for z = 0 µm and z = 60 µm. Right: T- field in the x-y plane. Acc. to [41]The iterative Rosenthal prediction method is applied to evaluate the influence of T- dependent variables on the prediction of thermal profiles. For this purpose, the determined fit functions for k and τ are inserted into the Rosenthal equation, which then is iterated. During the first iteration step, a high variation of melt pool dimensions occurs.

During the first iteration step, a high variation of melt pool dimensions occurs. After 5 to 10 iterations, the variation decreases, and after 15 iterations, the values can be assumed to maintain constant.To validate the iterative Rosenthal prediction, single tracks were processed. The width of the scan track is considered to indicate the width of the isothermal line where the first-order phase transition from liquid to solid takes place – marked in Fig. 3 as 𝑇𝑙. The width of the652processed single tracks based on the applied line energy density is plotted together with the predicted width in Fig. 4 for AMZ4.It can be seen that the analytically calculated melt pool width correlates to the measured scan track width. To determine the quality of the correlation, Fig. 5 displays the melt pool width in dependence on the energy line density [eq. 5].Figure 4: Single tracks of AMZ4. Left: Experimentally determined dimensions. Right: Analytically determined temperature fields.

5].Figure 4: Single tracks of AMZ4. Left: Experimentally determined dimensions. Right: Analytically determined temperature fields. Top: 𝑃𝐿 =20 W, 𝑣𝑠 =1 m/s. Mid: 60 W, 0.6 m/s. Bottom: 60 W, 1.6 m/s. [41]For AMZ4 the analytically determined values coincide with the measurement data up to a line energy density of 60 J/m. For higher energy input, a divergence of measured to calculated values can be seen. The basis for the Rosenthal approach is the consideration of conduction and diffusion as the main heat transfer mechanisms. This is assumed to be applicable for low energy input. At higher energy levels, mechanisms of evaporation, convection and keyhole formation can increasingly affect the heat flow and absorption. Thus, the deviation of measured to calculated values at higher 𝐸𝐿 can be explained.Figure 5: Experimentally and analytically determined scan track width. Left: AMZ4, 𝑃𝐿= const. Right: Vit101, varied 𝑃𝐿 and 𝑣𝑠. Acc.

Left: AMZ4, 𝑃𝐿= const. Right: Vit101, varied 𝑃𝐿 and 𝑣𝑠. Acc. to [40]To evaluate the influence of the process parameters on the thermal cycle, three different manufacturing parameter sets for the glass-forming alloys AMZ4 and Vit101 are examined. The conventional and the iterated approach are applied, and the results compared.653Cooling rates and temperature profiles in single tracksThe iterative Rosenthal approach allows a prediction of temperature fields in the x-y plane, which can be used to estimate the dimensions of the heat-affected zone (HAZ) implying different mechanisms of affection. These are melting, heating above Tg, and heating below Tg. Further, the heating and cooling rates within a single weld track can be calculated, which are meaningful for the vitrification. Fig. 6 (left) shows the t-T diagram for AMZ4 during one heating cycle. As shown in Fig.

Fig. 6 (left) shows the t-T diagram for AMZ4 during one heating cycle. As shown in Fig. 6, the cooling time is tc,AMZ4 = 1.3 𝑚𝑠 and corresponds to a cooling rate of R = 4∙10^5 K/s on the scan path at z = 0. This locates several magnitudes below the nominal critical cooling rate of 5 – 40 K/s for AMZ4 [28]. For Vit101, similar values are achieved.Figure 6: Left: t-T diagram for AMZ4 processed with 3 different parameter sets, comparing conventional and iterated model. Right: t-T diagram for Vit101 processed at 100 W and 0.8 m/s, at z = 0 µm and z = 50µm (lower layer in AM).To consider the effect of newly applied layers on the underlying material of the produced part, the re-heating of material at different heights of z = 0 µm and z = 50 µm (50 µm below the surface of the melt pool) is calculated and depicted in Fig. 6 (right) for Vit101. The iterated approach presents that at z = 50 µm the material is heated above Tg, but not molten.

6 (right) for Vit101. The iterated approach presents that at z = 50 µm the material is heated above Tg, but not molten. In this region, the concept of the cooling rate R needs to be adjusted since the entailed tc describes the cooling from the liquid state to the amorphous state. Instead, here the solid amorphous material is heated to a certain temperature and then cooled again. To describe this, the retention time tr is introduced as the time a certain point remains in the temperature regime between Tg and Tl:tr = 𝑡𝑇_𝑔𝑙𝑎𝑠𝑠_𝑡𝑟𝑎𝑛𝑠𝑖𝑡_𝑢𝑝 − 𝑡𝑇_𝑔𝑙𝑎𝑠𝑠_𝑡𝑟𝑎𝑛𝑠𝑖𝑡_𝑑𝑜𝑤𝑛Eq. 6The highest retention time is expected to exceed the highest cooling time and thus tr,max depicts a critical value for the margin of vitrification to partial crystallization. For Vit101 processed at 100W and 0.8 m/s, tr = 0.86 ms. For a first estimation, concepts concerning critical heating rates and their effect on vitrification issues are neglected here. If tr is inserted into eq.

For a first estimation, concepts concerning critical heating rates and their effect on vitrification issues are neglected here. If tr is inserted into eq. 3 in place of tc, it corresponds to a cooling rate of R = 5.56 ∙10^5 K/s. This is 5 magnitudes higher than the conservative Rcrit,Vit101 = 62 K/s. Accordingly, successful vitrification can be expected. Crystallization theoretically should not occur within single lines, even in the critical areas of the heat-affected zone.To discuss the influence of powder instead of solid material surrounding the heat source, the powder bed is approximated as solid with lower conductivity. In this case the thermal654conductivity can be expected to be about two orders of magnitude lower for powder materials compared to solid bulk material [42].The resulting cooling time according to Fig. 7 at x, y, z = 0 is then tc = 25 ms, with a resulting cooling rate: R= 520K / 0.025s = 20.8 ∙10^3K/s.

7 at x, y, z = 0 is then tc = 25 ms, with a resulting cooling rate: R= 520K / 0.025s = 20.8 ∙10^3K/s. The critical times of crystallization determined by Yang et al. lie between 14.2 and 22.4 ms for increasing oxygen content [43]. Thus, it can be concluded that at higher layer thicknesses, unsupported areas or downskin regions with an increasing ratio of powder to solid in the direct surroundings of the melt pool in combination with high oxygen contamination of the powder feedstock, crystallization could occur during PBF-LB/M manufacturing of AMZ4. Additionally, consecutive scan tracks within the same layer could affect the cooling times drastically. Due to the complexFigure 7: Logarithmic t-T diagram of AMZ4 comparing the proceeding of the temperature for powder and solid material, assuming k = 5.75 for solid material and k = 0.0575 for powder material at 293 K (no iteration).In theory, for single track manufacturing in PBF-LB/M no crystallization is expected for AMZ4 and Vit101.

The calculated cooling rates exceed the critical values by several magnitudes. In practice, partially crystallized samples were detected, which could be correlated to an accumulation of heat by subsequent scan vectors in x-y- or in z-direction (following layers). The pattern and sequence are part of the scan strategy, which can be customized. These findings motivate the consideration of BMGs with lower GFA for PBF-LB/M, such as the novel alloy Ti60Zr15Cu17S8.Transfer to feasibility of 𝐓𝐢𝟔𝟎𝐙𝐫𝟏𝟓𝐂𝐮𝟏𝟕𝐒𝟖In this section, the Rosenthal equation is applied to estimate the processability of the novel glass-forming alloy Ti60Zr15Cu17S8. Therefore, by Laser Flash measurement of cast material the thermal diffusivity and conductivity are determined in dependence on the temperature, as depicted in Fig. 8 (left). The conductivity shows an almost linear increase with rising temperature. The progression of the diffusivity shows a kink at a temperature of approximately 675 K.

8 (left). The conductivity shows an almost linear increase with rising temperature. The progression of the diffusivity shows a kink at a temperature of approximately 675 K. In general, the values for the second measurement run are higher than for the first run, which can be correlated to the effect of partial crystallization during the measurement. In order to estimate the general feasibility of vitrification, the t-T progression is calculated once at low temperature (295 𝐾; 𝜏𝐿𝑇=3.55 mm²/s; 𝑘𝐿𝑇= 8.76 W/ m∙ 𝐾) and once at655high temperature (1075 K; 𝜏𝐻𝑇= 5.27 mm²/s; 𝑘𝐻𝑇= 18.77 W/ m∙ 𝐾). The absorptivity was determined according to eq. 4 using cast material and is a = 0.31 for the chosen process parameters of P= 70 W and 𝑣𝑠= 1.0 m/s. The resulting t-T diagram is given in Fig. 8 (right).Figure 8: Left: Thermophysical properties 𝜏 and k of 𝑇𝑖60𝑍𝑟15𝐶𝑢17𝑆8.

The resulting t-T diagram is given in Fig. 8 (right).Figure 8: Left: Thermophysical properties 𝜏 and k of 𝑇𝑖60𝑍𝑟15𝐶𝑢17𝑆8. Right: t-T diagram derived by Rosenthal equation of 𝑇𝑖60𝑍𝑟15𝐶𝑢17𝑆8 at low (LT) and high (HT) temperature.For Ti60Zr15Cu17S8 the cooling time determined by Rosenthal is 𝑡𝑐= 0.66 ms, resulting in a cooling rate of R = 1.0 ∙10^6 K/s. For the chosen parameter set, both cooling rates are below the critical cooling rate. The melt pool length is 2 times higher at low temperature. This divergence leads to the assumption that for higher accuracy a fit function could be derived and applied to iterate the Rosenthal equation, as shown above for AMZ4 and Vit101. Nevertheless, this example shows, how the application of temperature-dependent thermal properties to the Rosenthal equation can enable an estimation of vitrification possibilities for novel materials and different processing parameters.

Thus Ti60Zr15Cu17S8 can be considered for manufacturing of BMGs via PBF-LB/M.ConclusionIn this work, the thermophysical properties of AMZ4, Vit101 and Ti60S were measured in dependence on the temperature. For AMZ4 and Vit101, fit functions were determined to describe the thermal conductivity and the thermal diffusivity. The absorption was determined for both materials by in-situ measurement. The resulting values were used to develop an iterative Rosenthal approach, considering the T-dependence of thermal conductivity and diffusivity. The approach was validated for line energy densities below 60 J/m (AMZ4) and 200 J/m (Vit101) and thus can be considered for future analytical applications. Divergence from experimental results is found for higher energy input, which can be explained by an increasing effect of keyhole mode and convection on the heat flow.

Divergence from experimental results is found for higher energy input, which can be explained by an increasing effect of keyhole mode and convection on the heat flow. The approach was then used to estimate cooling rates during the manufacturing by PBF-LB/M of the glass-forming materials AMZ4 and Vit101. The t-T diagrams along the z-height suggest, that the interface between melt pool and heat affected zone is most likely to experience a crystallization, since here the thermal cycle is prolonged. Although the iteratively determined cooling rates exceeded the values according to conventional Rosenthal approach, the nominal Rkrit is met for all three alloys.For the novel glass-forming alloy Ti60Zr15Cu17S8, cooling rates and melt pool dimensions were calculated via Rosenthal, considering temperature-dependent thermal conductivity and diffusivity. The resulting cooling times match the allowance.

The resulting cooling times match the allowance. Thus, this656material class is expected to be a promising candidate for BMG processing via PBF-LB/M. The influence of heat accumulation due to subsequent scan tracks and layers needs to be considered in future experiments and simulations to ensure the vitrification on larger scales.It is reported that single lines of all 3 investigated materials can be processed with cooling rates high enough to enable vitrification.

For AMZ4 the limits of processability are highlighted: When heat accumulation (e.g., due to surrounding powder material) occurs combined with higher oxygen contamination of the powder feedstock, the critical cooling times can be exceeded, indicating a risk of partial crystallization.The results can be summarized as:1) Within single tracks, successful vitrification is theoretically expected.2) High cooling rates should enable processing of BMGs with lower GFA.3) Possible limitations for vitrification: Scan strategy, oxygen contamination in powder feedstock, lack of heat transfer via solid / support.OutlookFurther optimization steps for increased accuracy of the Rosenthal equation could be performed. It can be suggested to investigate the influence of heat convection on the heat flow, enabling the verified use of an analytical approach to estimate cooling rates and melt pool dimensions at higher energy inputs.

Further, material parameters such as 𝑇𝑔, 𝑇𝑥 and 𝑅𝑐𝑟𝑖𝑡 not only depend on the temperature but also the cooling/heating rates, and therefore a time- dependence could be considered. The quality of raw material also highly affects 𝑅𝑐𝑟𝑖𝑡 and could be considered an important factor for quality control as well as for comparison on the side of the limitations of processability.Additionally, other analytical approaches could be of interest to evaluate and optimize the predictions. Goldak developed an approach based on assuming an elliptical heat source in welding, eliminating the issues concerning point heat sources in the Rosenthal equation [44]. This could be used for likewise simulation, verification, and improvement of existing models.AcknowledgementsThe authors would like to thank Prof. Roland Schmechel from the Institute of Technology for Nanostructures at the University of Duisburg-Essen for their support in carrying out Laser Flash measurements.

Roland Schmechel from the Institute of Technology for Nanostructures at the University of Duisburg-Essen for their support in carrying out Laser Flash measurements. This research was supported by the German Research Foundation (DFG) through Grant No. KL 3357/1-1. The PBF-LB/M system SLM 280 Hl was funded by the German Research Foundation (DFG) through Grant No. INST 20876/341-1.657References [1] T. Wohlers, R. I. Campbell, O. Diegel, J. Kowen, N. Mostow, and I. Fidan, Wohlers report 2022: 3Dprinting and additive manufacturing : global state of the industry. Fort Collins, Colo., Washington,DC: Wohlers Associates; ASTM International, 2022.[2] S. Pauly et al., “Processing metallic glasses by selective laser melting,” Materials Today, vol. 16,1-2, pp. 37–41, 2013, doi: 10.1016/j.mattod.2013.01.018.[3] N. Sohrabi, J. Jhabvala, and R. E. Logé, “Additive Manufacturing of Bulk Metallic Glasses— Process, Challenges and Properties: A Review,” Metals, vol. 11, no. 8, p.

alyzing Product Lifecycle Costs for a Better Understanding of Cost Drivers in Additive Manufacturing C. Lindemann*, U. Jahnke*, M. Moi*, R. Koch*Direct Manufacturing Research Center (DMRC) and Chair of Computer Application and Integration in Design and Planning (C.I.K.), The University of Paderborn, Mersinweg 3, 33098 Paderborn, GermanyAbstractThe costs of additive manufactured parts often seem too high in comparison to those of traditionally manufactured parts, as the information about major cost drivers, especially for additive manufactured metal parts, is weak. Therefore, a lifecycle analysis of additive manufactured parts is needed to understand and rate the cost drivers that act as the largest contributors to unit costs, and to provide a focus for future cost reduction activities for the Additive Manufacturing (AM) technology.

A better understanding of the cost structure will help to compare the AM costs with the opportunity costs of the classical manufacturing technologies and will make it easier to justify the use of AM manufactured parts. This paper will present work in progress and methodology based on a sample investigated with business process analysis / simulation and activity based costing. In addition, cost drivers associated with metal AM process will be rated.IntroductionNowadays, industrial companies face more and more complex challenges in product development. Customers ask for innovative, individually tailored products with a high product quality for a reasonable price. In addition, the economic lifespan of products decreases which forces the companies to shorten their time to market and their development cycles [SBA02]. Through globalization the competition in fertile markets increases.

Through globalization the competition in fertile markets increases. Imitators from foreign markets make it harder for the companies to maintain achieved market shares [Bu94]. One solution to increase innovation and shorten the time to market is delivered by a new production technology: Additive Manufacturing. AM, also known as Rapid Manufacturing, is a technology development of Rapid Prototyping, which was established around 1986 mainly based on Stereolithography. Nowadays there exist several different technologies for AM. They all have in common that they are based on 3-D product data and that they manufacture layer by layer. The difference between AM and Rapid Prototyping is based on the product characteristics. While prototypes are used to show special product properties or functions during the product development phase, AM delivers parts with characteristics of a final product [Geb07].

While prototypes are used to show special product properties or functions during the product development phase, AM delivers parts with characteristics of a final product [Geb07]. The additive production process allows product designers to create parts with high geometric freedom. Nearly all shapes can be realized by AM, which consequently allows the designer to improve product properties [Geb07]. This groundbreaking technology has the potential to revolutionize the theory of product design. It allows the designers to concentrate on product features instead of manufacturing restrictions given by the existing production technologies. Unfortunately, the knowledge about this technology and the freedom of design is not widely spread among today’s engineers [BLR09]. Thus, the potentialities of AM often go unused and the added value compared to traditional manufacturing remains low.

Thus, the potentialities of AM often go unused and the added value compared to traditional manufacturing remains low. Particularly branches dealing with long product life cycles could benefit from weight savings and the integration of several functions in one part without the need to assemble. Within a study [EKW+12a], especially the aircraft production, automotive production and the electronics industry have177been identified as very promising to profit from the use of AM. Experts have selected these branches by means of assessing the prospective attractiveness of current application fields [EKW+12a]. But nowadays the process costs of AM are one of the top three critical success factors of this technology [Gaus09]. This seems to be one of the major barriers for the further dissemination of this technology.

This seems to be one of the major barriers for the further dissemination of this technology. In order to promote the dissemination of the AM technology the economic aspects and the costs need to be evaluated in detail.Economic Aspects of Additive ManufacturingThe fast moving markets require a rethink in manufacturing. As products are getting more individual, and at the same time, getting a shorter lifecycle, AM enables the manufactures to fulfill these requirements.Design / Construction Compared to traditional manufacturing, the general advantages of AM are the capabilities in design and development of products. Despite certain limitations, companies are using AM increasingly to use the many possible benefits like complexity-for-free manufacturing. In traditional manufacturing there exists a direct connection between complexity and manufacturing costs. A relationship tying cost to complexity does not exist in AM.

In traditional manufacturing there exists a direct connection between complexity and manufacturing costs. A relationship tying cost to complexity does not exist in AM. On the basis of the procedure there’s almost no limitation in relation to the complexity of geometry and this without the need to produce any kind of tools (e.g. forming tools). This is why one speaks of “complexity-for-free” [HHD06, p. 9]. Consequently, most restrictions of design for manufacture and assembly are not valid for AM. Designs intended for traditional manufacturing are often heavily limited by high costs in construction and tool-making. The greater freedom of design via AM makes it possible to combine an assembly of parts into one part and, therefore, to reduce the required assembly work and costs. In addition, no compromises regarding the assembly capabilities are necessary [HHD06, p.

In addition, no compromises regarding the assembly capabilities are necessary [HHD06, p. 5].Capabilities of Additive Manufacturing As mentioned above, AM makes it possible to replace several traditional manufactured and assembled parts by one part. Hence, this allows an integration of functions from different parts, which may result in better performance. Even if the requirement on movability of a part in relation to standing parts exists, e.g. ball and socket joint, the production with AM can be done completely as a single monolithic structure. Unlike the applied design rules for traditional manufactured parts, which do not apply to parts produced by AM, design guides for AM processes must be considered. Due to the freedom during the design, the simple assembly moves forward into focus and becomes of higher importance. The fewer number of parts and fewer assembly steps may result in a high impact to production costs.

The fewer number of parts and fewer assembly steps may result in a high impact to production costs. The targeted design of a relieved or decreased assembly may result in a much higher reduction of the to parts designed for traditional production costs than manufacturing [BDK94]. A lesser amount of parts provides other advantages as well. For example, fewer parts must be sourced, labeled and evaluated. In the end this also reduces the spare parts that have to be stocked. Since there is no need of tooling for production of spare parts, it is unnecessary to hold legacy tooling in storage. The complexity of the production and the whole management decreases and therefore savings in the entire business can be achieved [GRS10, p. 288f].

The complexity of the production and the whole management decreases and therefore savings in the entire business can be achieved [GRS10, p. 288f]. As a consequence of the simple way to produce complex geometric structures, the offer of mass-produced individual parts, like it is done for hearing aids, is given.the construction compared178Economic application There are key economic criteria in the area of AM, which cannot be directly expressed monetarily. AM helps to shorten the time-to-market-duration. The resulting advantage is not allocable by hard facts in general [Zaeh06, S. 121]. Another opportunity of AM is the increase of diversity of variants, while quantity of variants decreases. At this point, AM technologies can play out their potential because they precisely support the individualization. Additionally, the development is much faster which makes early market positioning possible. The discontinuation of the need for production tooling is another key factor.

Additionally, the development is much faster which makes early market positioning possible. The discontinuation of the need for production tooling is another key factor. Further, the introduction of new products is much less risky than before, due to the elimination of costly production tooling. This has a strong impact on the post-processing of existing products. Changes in the design can be published to the market even faster. With a slight or agile production significant improvements of the operational efficiency are achievable, therefore, the usage of such a technology is a competitive advantage. [Zaeh06, p. 121f] [HHD06, p. 160f] However, it should be noted that customers might not be prepared for higher prices or loss of quality for individualized products. Therefore, new technologies must enable a flexible and inexpensive production.

Table 1 summarizes the pros and cons in usage of AM during product lifecycle management.Table 1: Pros and cons in product lifecycle management [Zaeh06][HHD06][GRS10]ProsMore flexible development - Freedom of design and construction - - Less assembly - No production tools necessary - Less spare parts in stock - Less complexity in business because of less parts to manageIntegration of functionsCons - Available software is a limiting factor - High machine and material costs - High calibration effort - Quality of parts is in need of improvement - Rework of parts is often necessary(support structures)Building time depends on the height of the part in the building chamberNo tools for productions need to hold instock (only digital/CAD data) - Less time-to-market for products - Faster deployment of changes - Offer of individual productsAdditive manufacturing for supply chain management As mentioned above, with the ability to produce highly complex parts without tools, a decrease of production costs is possible.

Since there is no need to produce a high amount of an individual part to refinance the tools, like in traditional manufacturing, AM is predestinated for low volume production. Hence, affordable and high complexity individual products can be manufactured. To achieve a holistic economical understanding further aspects must be considered. This includes e.g. expenses, logistics, supply chain and other variables. AM with its unique characteristics establishes a complete new business model compared to business models for traditional companies using traditional methods to manufacture [HHD06, p. 159f].Methodology for accessing lifecycle costsIn this paper we use the definition of the intrinsic lifecycle. Intrinsic lifecycle means all different steps a product passes, from the first product concept trough the production processes and ending with the disposal of the product [PaBe07]. These processes will be modeled and investigated for lifecycle costs.

These processes will be modeled and investigated for lifecycle costs. "Lifecycle costs are the costs associated with179product users as the sum of all costs due to the purchase and during the period of use of a product (plant, machinery, equipment, apparatus) (product life)" [EKL07]AM is capable of cutting costs in a variety of areas. This starts with the reduction of parts and, therefore, a saving of time required to design a part. Tests and tolerances can be reduced in a significant way [DiCu11] as well as warehousing and assembly costs. The integration of an assembly into a single part also means that fewer subcontractors need to be controlled and managed. Digital tooth cap (invisalign®) production has demonstrated the ability of optimizing the whole process chain. The transition from analogue global dental production to digital local production has saved 85% of the logistic steps, reducing the production as well as the energy consumption for production activity by 80 %1.

But the ability to realize cost-savings and added value does not stop in only the company but also benefits the customer in increased functionality and reduced lifecycle cost for the usage of a product. Redesigns of sample aerospace parts have shown a weight reduction potential of up to 70 % of the original part weight. If the weight of an aircraft is reduced by 1 kg a year this saves $3000 in fuel annually [Wes11], keeping in mind that an average aircraft has a thirty years’ lifespan. This means especially for this sector only the weight savings have an incredible capability of reducing lifecycle costs. But not only weight reduction can benefit the costs of a product. Considering all these aspects, it is obvious that it is hard to compare the costs of a product simply based upon the production costs per piece. Therefore, a lifecycle based approach needs to be taken.

Therefore, a lifecycle based approach needs to be taken. The goal of the CoA²MPLy2 project is to understand and rate the cost drivers that act as the largest contributors to unit costs and to provide a focus for future cost reduction activities for the AM technology. This will help to expand the applications for additive manufactured parts focusing on Metal Additive Manufacturing (MAM). A better understanding of the cost structure will help to compare the AM costs with costs of the classical manufacturing technologies and make it easier to justify the use of AM technology. This project focuses as well on the costs over the product lifecycle, starting in the design phase and ending with the spare part supply, which is similar for all AM technologies.

This project focuses as well on the costs over the product lifecycle, starting in the design phase and ending with the spare part supply, which is similar for all AM technologies. By identifying the cost structure, one will be able to understand how much of an economic impact ‘design optimization’ of part geometry has on unit costs.The first objective is to identify the machine rate cost structure relative to current state in AM. The second objective is to understand how conventional machining compares to the MAM machine rate unit cost structure. Therefore, current business processes will be analyzed. An exemplary part will be redesigned to demonstrate the advantage of AM to conventionally machined part costing and assembly. A list of the major cost drivers will be generated to provide guidance, in which reduction will become areas of opportunities for AM. Currently the project is in phase 2 (compare figure 1).

A list of the major cost drivers will be generated to provide guidance, in which reduction will become areas of opportunities for AM. Currently the project is in phase 2 (compare figure 1). Future work will be based on the results presented.technology andto comparefigure 1: Methodological approachcost1www.melotte.be 2 CoA²MPLy: “Costing Analysis for Additive Manufacturing during Product Lifecycle”180itCosting Models for ProductionTo date few approaches have been taken to calculate the production costs of AM. Most of them have been looking at the Laser Sintering (LS) and Fused Deposition Modeling (FDM) processes and have compared them to injection molding or similar. The most important will be presented in this section. Some of the most influencing authors were Augsburg, Hopkins/Dickens (HD), Ruffo/Tuck/Hague (RTH) and Gibson/Rosen/Stucker (GRS).

The most important will be presented in this section. Some of the most influencing authors were Augsburg, Hopkins/Dickens (HD), Ruffo/Tuck/Hague (RTH) and Gibson/Rosen/Stucker (GRS). A comparison of the models shows that all authors have chosen a similar approach for the calculation of costs in their models. Each of them has set a specific emphasis on a certain topic. The model of Augsburg for example describes manual process steps (e.g. removal of support structure, machine preparation), which helps to assign the labor costs directly to the product. The model of RTH is the only model showing the composition of the indirect costs and the composition of the cost rate for the build time. The Model of GRS shows advantages in the estimation of realistic building times, as no other model is as detailed in this section, as well as in the estimation of the powder usage for the production process.

In summary, one could say that each of the existing costing models has advantages and disadvantages, but no model meets all criteria satisfactory. Thus, there is a need to combine the strengths of the existing models without including their weaknesses and to develop a new costing model that is suitable for the calculation of today’s AM. A model considering the whole production process completely and correctly will be the basis for research how to include for instance development and lifecycle costs.Development of a costing modelPrior to the development of a cost model, all cost relevant processes of the AM production process have been investigated and modeled with Event Driven Process Chains (EPC).figure 2: rough Production model181Business process models, especially EPC’s, have been proven in the past that they are capable to represent knowledge. They are the basis for an aggregation of different types of information [LPK10].

They are the basis for an aggregation of different types of information [LPK10]. This should help to understand the processes in detail. Important process data has been captured. On this basis simulations of the production process have been performed in order to estimate the different influences of the cost drivers of the model. This helped to simplify different processes, as a cost model may not become to complex.As a calculation method a “Time driven Activity Based Costing” (TD-ABC) approach has been taken. This approach allows the consideration of different influence factors on the basis the use of resources [CFG07].

This approach allows the consideration of different influence factors on the basis the use of resources [CFG07]. For the estimation of cost relevant processes, the process steps of the initial model have been simplified into four main processes:Preparation of the building job • Production of the building job • Manual removing of sample parts and support • Post processing to enhance material propertiesThe main processes have been selected in order to be able to represent different cost centers. This facilitates the calculation and makes it easier to adopt the model to different production environments.The costing center “preparation” includes all steps from the initial CAD data, placement of the parts in the building chamber and creating the support structures. Interviews with engineers have shown that the costs in this step are highly dependent on the complexity of the different parts and the complexity of the building job itself.

Interviews with engineers have shown that the costs in this step are highly dependent on the complexity of the different parts and the complexity of the building job itself. Therefore, a complexity factor has been introduced in the calculation of this cost center, which is capable of estimating the duration of this task. This is especially important, as this process requires an experienced engineer, whose labor costs represent the main costing factor in this step. Furthermore, this factor allows distinguishing between the variation of an existing building job as well as the creation of a new one.The costing center “machine” can be seen in detail in figure 3 and is mainly based on the work of RTH and Dietrich [Diet10]. This production model is completely time driven and gets an additional fixed cost rate for the initial gas flooding of the machine as well as a fixed labor cost rate, as the processes are considered constant for each building process.

The hourly machine rate consists of an indirect and a direct cost rate and is multiplied with the build time afterwards. Many factors are already available as the energy consumption of different machines (compare[BTL+11]).CostsBuild = CostsFixed + MachineHourlyRate * BuildtimeThe estimation of the build time is certainly one of the most important factors in costing models as it influences the costs of the build significantly. A lot of authors have put some effort in the calculation of the building time (compare [MeRe11] and [GRS10]. Therefore, different approaches can be taken. As stated above the model of GRS is the most detailed concerning the estimation of the build time. Hence, the equations from GRS have been adopted for that model. Further research attempts to find a way to simplify the estimation of the building time depending on factors such as standard elements, material selection, part density and part alignment.

Further research attempts to find a way to simplify the estimation of the building time depending on factors such as standard elements, material selection, part density and part alignment. As material usage is not directly related to the build time these costs will be allocated in an additional step.182Machine	
 Costs per	
 BuildFixed	
 CostsLabour	
 CostsBuild	
 TimeGas	
 CostsMachine Hourly	
 RateOverall Working	
 TimeIndirect Cost_rateDirect	
 Cost	
 RateMaintenanceOther	
 CostsProduction OverheadMachine PurchaseDepreciation TimeElectrical Engergy 	
 per	
 HourGas	
 Costs per	
 HourHardware CostsDepreciation TimeSupporting EquipmentDepreciation TimeSoftwareDepreciation TimeFacility RentDepreciation TimeLabour totalLabour totalfigure 3: Machine Costs per buildAfter the building process, the costing model allows different treatments to enhance the mechanical properties of the build.

This process step can be considered additionally because it depends strongly on the product requirements if a pre treatment is necessary. As heat relief treatments or hot isostatic pressing (HIP) have proven to improve material properties significantly this was found useful by the authors to estimate total part cost for AM. An hourly machine rate, which consists of direct and indirect costs, is multiplied by the processing time.The last costing factor sums up all necessary manual processing steps. These include steps like support removal, polishing and quality control. An average value for the processing time has been taken which is multiplied with the labor costs per hour. These can be estimated lower than in the preparation phase because only a minimally skilled worker is needed to perform the tasks.

These can be estimated lower than in the preparation phase because only a minimally skilled worker is needed to perform the tasks. The whole processing time will be multiplied with a complexity factor capturing the different efforts that it takes to remove support structures.As the material costs have not been part of the machine costs per build, they need to be added to the costs of the other cost centers. The formula to calculate the material costs is the same as used by GRS as it is suitable for the FDM, SLM and LS processes.MaterialCost = Supportstucture ∗ Wastefactor ∗ Number of Parts ∗ Partsvolume∗ MaterialPrice ∗ MassDensity183Calculation of a sample part and interpretation of resultsAs more research and parts are needed for a generalization of the results, this paper will discuss AM-production-costs on the basis of a sample part. For the calculation one of the sample parts used by Augsburg will be benchmarked.

For the calculation one of the sample parts used by Augsburg will be benchmarked. Its initial use can be found in the automotive industry. As a first evaluation of the costing model the cost of this sample metal part has been calculated. The considered processes started with the data preparation followed by manufacturing and post processing. The sample part has retrieved an additional heat treatment afterwards to increase mechanical properties.Figure 4 illustrates the cost breakdown for the sample part calculated with the aforementioned model.

Some major parameters for the Model were:AM-machine utilization: 4500 h/year • Depreciation time: 5 years • Investment costs: 500,000€ • Costs for maintenance 21,666 €/ Year • Build rate: 6.3 cm³/h • Build Material: Stainless Steel 316L • Material Price 89 €/kg • Part Volume: 1cm3 • Layer thickness 0.3 µm 4,25%2,62%1,37%Costs	
 for	
 PreparationBuildingprocess_>ix11,37%Oven6,65%73,73%Material	
 CostsPostprocessingMachine	
 Costsfigure 4: Sample Part from Augsburg and the given cost breakdown structure by a batch size of 190 parts per build184Costs	
 per	
 piece150,00	
 €100,00	
 €50,00	
 €Costs	
 per	
 piece0,00	
 €0100200300400500600figure 5: Costing Curve depending on numbers of the buildThe major cost driver consists mainly of machine costs (73%), as other authors have stated, followed by the material costs, which only make 12% of the total costs.

The post processing process, followed with a similar amount, by the preparation process, represents the third largest cost driver. In figure 5 one can see the relationship between costs per unit and the number of manufactured parts. The star indicates the production quantity, which is used for the following part of the paper. You can see that the initial costs for a single piece rapidly drop as more parts are being placed in the building chamber. The developed costing model does not show the typical chainsaw effect, which can be seen in [RTH05] for example. A slight effect can be noticed after a new production process has been initiated and the building chamber is not fully utilized. Over the time, this effect minimizes as the costs are split on more parts. That means after a certain amount of produced parts the influence of a fully utilization of the building chamber decreases.

Over the time, this effect minimizes as the costs are split on more parts. That means after a certain amount of produced parts the influence of a fully utilization of the building chamber decreases. It has to be stated that the costs of the build are not related to the complexity of the part itself as stated in [HHD06] - meaning that the utilization degree of the building chamber has a small effect on the costs in mass production. As the production capacity increases, the curve converges to a straight line. Figure 6 shows variation of some of the above stated cost drivers.figure 6: Variation on different Influence Factors in %185Therefore, for each column one cost driver has been varied in order to see the effects on the composition of the total costs. Figure 7 shows the total distribution on the costs of the build.figure 7: Variation on different Influence factors in €Discussion of ResultsThe largest contributor for building costs are the machine costs.

The variation of influence factors have shown that a reduced machine rate cost can be achieved but will stay one of the dominant factors in the production process. One reason is that the labor costs for the AM building process itself can be reduced to the loading and unloading of the machine, as the production process is a blind process. As the material costs are not considered as a part of the building process, these cannot contribute to a decrease in machine rate costs. As the process is a fully automated and “lights-out” process it is logical that the machine rate costs have the greatest contribution to the total costs of a build. As GRS states, the changing allocation of the overhead costs to the production enhances this effects.The material costs as the second largest cost driver has a certain influence on the building costs as well. For the sample part this is smaller than for a single piece high volume build.

For the sample part this is smaller than for a single piece high volume build. In general, the volume of additively manufactured parts will decrease constantly as the designer is able to construct independently of manufacturing restrictions. Lattice structures as researched in [GLJ+11] have the potential to reduce material volume and therefore the cost of the build. Materials like titanium are still very expensive and can raise costs up to nearly 50% even for low volume parts. As more material is sold every year [WOH12], more manufacturers will enter the market and the costs for the material, which in some cases (e.g. titanium), are approximately ten times more expensive than traditional materials, will decrease in the future. Therefore, the influence of the chosen material on the total cost, especially for low volume parts will decrease even further in the future.These two cost drivers are followed by the post and pre processing of the parts.

While the production process itself is nearly labor-free, the post processing is not yet automated. In fact,186the only two processes in which labor costs are significant, are the data preparation as well as the post processing, because it is necessary to remove the support structures etc. The direct labor costs for the production process are locked in the cost driver “Buildingprocess fix” which amount to a very small portion of the total part cost. The third largest cost driver is represented by the costs for the data preparation. This influence may be greater than the expected, since only a simple building job has been investigated. For the preparation process a skilled and experienced engineer is necessary. Thus, the main cost driver of the data preparation process is represented by labor costs. Knowledge is necessary to place the parts in the building chamber.

Thus, the main cost driver of the data preparation process is represented by labor costs. Knowledge is necessary to place the parts in the building chamber. Further research will bring more knowledge about the placing of parts and new software will automate this process so that a computer-based placement of parts in the building chamber will reduce the cost driver and make the technology more viable. As the main factor for the preparation is labor costs to place the parts in the building chamber, it only has to be performed once for larger series of parts. This will decrease the costs for the data preparation. Thus, the developed cost model allows a differentiation between mass produced parts (only place parts once) and customized combination of parts in the building chamber.

Thus, the developed cost model allows a differentiation between mass produced parts (only place parts once) and customized combination of parts in the building chamber. An automated placing of parts in the building chamber on the other hand would make this differentiation obsolete.In the calculations the main adjustment “knobs” for influencing the cost-driver of the machine rate costs were identified as the working load (or “degree of machine utilization”) and the building speed of the machine. New technologies, such as the use of a dual laser concept by SLM3 will increase building speed in the future. A standardized way to determine the building speed of the machine in combination of different materials needs to be found in order to compare building rates of different processes. The results show that AM is mainly attractive in terms of additive batch production for suppliers or companies who can reach a high degree of machine utilization.

The results show that AM is mainly attractive in terms of additive batch production for suppliers or companies who can reach a high degree of machine utilization. Therefore research should focus on the enhancement of maximizing the degree of utilization as well as on the increase of the build-rates of the machine.Figure 8 shows the comparison of the initial and realistic value setting for the costing calculation and an optimized future production.

The following assumptions have been applied to the original costing model:Assumptions:Building rate: 20 cm3 • Utilization rate: 7800 h/year (that means 90%)Material Costs: 40 €/kg (for the same material)Machine investment costs: 320,000 € As you can see in the results the combination of different factors may reduce part costs by over 50% in comparison to the actual market price.figure 8: Change of Values3http://www.slm-solutions.com/de/produkte/slm-anlagen/slm-280-hl/187Conclusion and OutlookTo estimate the benefits of AM a lifecycle-based approach has to be taken into consideration. Otherwise the advantages of this technology will be easily underestimated. The costs of AM builds are very complex but are not the only costs that have to be considered. A first step was taken by modeling the production process. It seems more efficient to work on technological aspects than to only reduce costs of material and the purchase price of the machine.

A first step was taken by modeling the production process. It seems more efficient to work on technological aspects than to only reduce costs of material and the purchase price of the machine. Further research follows, based on the comparison of an optimized aerospace sample part. These will include the cost estimation based on different standard elements. The rating of the cost drivers so far has shown that there still is a significant cost reduction potential. Still labor costs make a significant part on pre and post processing of a build. The existing process model will be enhanced by lifecycle processes. Looking for example at the aerospace industry weight reduction provide the possibility to save the customer a lot of expenses during the usage of the product through lightweight constructions. As this example already shows, the costs and benefits of AM strongly depend on the industry of usage. Also quality assurance costs have to be taken into account.

As this example already shows, the costs and benefits of AM strongly depend on the industry of usage. Also quality assurance costs have to be taken into account. These are significantly higher in the aerospace or medical industry compared to other industries. That highlights the need of an industry-specific investigation of AM-costs over the whole lifecycle. The machine utilization rate will be replaced by in the future with the Overall Equipment Effectiveness as a superior, more accurate way to gauge system performance based on customer demands. Influence of the designer on lifecycle and production costs will be part of the investigation as well.Literature[BDK94] [BLR09] [BTL+11][Bu94] [CFG07][DiCu11][Die10][EKL07] [EKW+12a][EKW+12b][Gaus11][Geb07][GLJ+11][GRS10][HHD06][LPK10][PaBe07][SBA02][WES11] [Woh12]Boothroyd, G.; Dewhurst, Peter; Knight, W. A. (1994): Product design for manufacture and assembly. New York: M. Dekker.

An in-situ MPM image captures the independent melting condition at specific positions within each build layer. However, identifying material defects such as horizontal lack-of- fusion using individual, discontinuous MPM images poses multiple challenges. This paper builds upon the authors’ previous work on data registration and data fusion to analyze material remelting conditions based on co-axial MPM images. MPM data are fused to formulate layerwise remelting ‘maps’ to evaluate printing quality. A 3D part with various overhang features was built to verify the method. A regular layer with solid support from the previous layers has around 30 % remelting ratio for the given laser scan conditions. In contrast, overhang regions remelted at about 10 % with the same process setting, whichis too low to provide sufficient material fusion, resulting in lack-of-fusion between melting tracks.

In contrast, overhang regions remelted at about 10 % with the same process setting, whichis too low to provide sufficient material fusion, resulting in lack-of-fusion between melting tracks. The negative impact in remelting would not immediately disappear in subsequent layers following the overhang regions. Results shows three additional layers are required to fully recover the remelting condition back to normal. The remelting results from MPM are also visible within layerwise optical images of the same surface.IntroductionLaser powder bed fusion (L-PBF) process uses laser to melt the powder spread on the build chamber. The material supplying mode in part formation is different than those additive manufacturing (AM) machines with a nozzle such as laser direct energy deposition (L-DED).

The material supplying mode in part formation is different than those additive manufacturing (AM) machines with a nozzle such as laser direct energy deposition (L-DED). Research shows the L-PBF can create parts of higher strength and ductility than L-DED [1], the random placement of particles in the powder bed itself introduces stochastic variability during the melting-solidification process [2]. In addition, variations in the laser beam, scan path, and build chamber environmental conditions introduce randomness that may cause irregular melt pool formation and lead to over-melting or under-melting [3, 4]. Complex geometric features add more difficulties in controlling melt-pool formation [5, 6].

Complex geometric features add more difficulties in controlling melt-pool formation [5, 6]. However, defects associated with complex geometric features and process settings can potentially be predicted at the design stage, and, hence, they may be avoided during the process through re-design of process settings, or model-based feedforward control [6-8].It is essential to understand the correlations between melting conditions and complex geometric features for defect prediction. For clarity, two concepts need to be explained – melting and remelting. Melting refers to the powder material initially changing from solid phase to liquid phase by receiving thermal energy from the laser beam [9]. The instantaneous molten region forms the melt pool, which traverses over specified areas of the powder layer using a prescribed scan strategy. The area over a powder layer that is traversed by the melt pool defines the solid portion of the AM part within that layer.

The area over a powder layer that is traversed by the melt pool defines the solid portion of the AM part within that layer. It is critical to notice that the laser melts not only powders but also remelts previously solidified materials—either adjacent within the same layer, or remelts portions of the lower solidified layers. Remelting, as the name suggested, represents melting those previously solidified melt pools [10]. Remelting typically creates the overlapping region between adjacent melting tracks and establishes the consolidation of material within AM parts [11]. Defects, particularly lack-of-fusion porosity,339may occur when there is insufficient remelting, and is known to be a function of the melt pool geometry, the scan strategy, and the powder layer thickness [12].

On the other hand, excessive remelting may introduce different types of defects such as keyhole porosity, or heterogeneous microstructure [13, 14].Figure 1 shows examples of different remelting conditions, conceptually. The figures are the cross- sections of two melting tracks. The dashed lines mark the laser beam center. The blue patches are remelting areas, and the orange circles are un-melted powders. The hatch spacing is the distance between the two tracks. Figure 1(a) demonstrates the normal remelting condition. The melt pool tracks have the ideal size and position to create just enough overlapping area between tracks and layers [15] leaving no un-melted powder. The resulting part would be fully dense once completely solidified. Figure 1(b) shows the under-melting condition while the melt pool itself has no initial defect. Insufficient remelting could be caused by poor scan strategy design, such as too wide hatch spacing, or layers that are too thick.

Insufficient remelting could be caused by poor scan strategy design, such as too wide hatch spacing, or layers that are too thick. Nevertheless, the under-melting in Figure 1(c) is mainly the result of irregular melt pools, or varying surface topography of the lower layer. The melt pool in Track 2 is significantly under-sized. As a result, it cannot establish enough overlap with Track 1, which has regular-sized melt pools. The small remelting area left un-melted powders around Track 1. These voids would form porosity in the final part. The causality of condition in (c) can be varied. On the other hand, oversized melt pool overlapping multiple times can cause over-melting condition.

These voids would form porosity in the final part. The causality of condition in (c) can be varied. On the other hand, oversized melt pool overlapping multiple times can cause over-melting condition. Common reasons for irregular melt pool formation are improper energy input that results in unstable melt pool formation (e.g., balling) [16], systematic fluctuations in local heating as a function of the scan strategy [7], or variations in local thermal conduction due to complex geometric features such as overhangs [8]. This paper chiefly focuses on the last condition—the impact caused by overhanging features.Figure 1. Conceptual remelting conditions between two melt pool tracks. Note, realistic melt pool geometries could be different. Blue patches represent remelting regions and orange circles are raw powder. (a) Proper melt pool size and hatch spacing create ideal remelting. (b) Abnormal remelting can happen while melt pool size cannot fill designed hatch spacing.

(a) Proper melt pool size and hatch spacing create ideal remelting. (b) Abnormal remelting can happen while melt pool size cannot fill designed hatch spacing. (c) Irregular melt pool can cause abnormal remelting.The overhanging surface is also called a ‘downward-facing’ surface or ‘down-skin’ surface [5, 17]. Instead of being supported by a solid layer, the overhanging surface is supported by powder. Most of the overhanging surfaces are created with purpose and specified by part geometry and build direction. Figure 2 shows an example of a typical overhanging surface. The cross-section of the part is on the left, where the dashed curve340is the overhanging surface. According to the layer-by-layer building process in L-PBF, the amplified view of the overhanging surface shows that the tip of each layer would always build on un-melted powders. The loose powder does not have the same strength as the dense solid.

The loose powder does not have the same strength as the dense solid. As a result, research shows that the overhanging region can exhibit defects such as craggy surface (or ‘dross’), lower density, and inferior geometric deviation [5, 17].Figure 2. Overhang impact to powder bed fusion process.Existing studies on overhang geometries built with AM are diverse in scope and methodology [17-20]. Simulations using finite element analysis or computational fluid dynamics mainly investigate the effect of heat accumulation in the vicinity of the melt pool due to the lower thermal conductivity of the powder compared to solid. Changes in surrounding material lead to locally variable thermal heat diffusion, and are prone to irregular melt pools on the overhanging region [5]. Experimental methods typically use traditional optical or mechanical measurement techniques to characterize AM overhang features.

Experimental methods typically use traditional optical or mechanical measurement techniques to characterize AM overhang features. For example, scanning electron microscope (SEM) images show that residual pores are widely distributed in the overhanging region [5]. The rough overhanging down-skin surface has been characterized in various AM parts [21]. Those critical findings gathered evidence of the defects associated with overhang features. However, few studies examine what happens precisely during the build process or the physical conditions that lead to those defects. High-quality in-situ monitoring data may explain the transient states of the build process through advanced data analytics, which could undercover the causal relationships between overhang geometry, variable thermal conditions around the melt pool, and overhang defects [22].This paper aims to analyze the remelting condition during the overhang build process based on high frequency in-situ monitoring data.

This work designs a specific experiment that collects millions of in-situ monitoring images of a part with overhang features. We deploy the previously developed AM data registration [23, 24] and AM data fusion techniques [11] to reconstruct the remelting maps of L-PBF layers. The last two sections present and discuss the statistical and visualization results of the remelting conditions on overhang features.Experimental DesignThe Additive Manufacturing Metrology Testbed (AMMT) at the National Institute of Standards and Technology (NIST) is the primary platform developed to study powder bed fusion processes. AMMT is a fully customized metrology instrument that enables flexible control and measurement of the L-PBF process [25]. It is equipped with the capability of precise laser beam control and high-rate in-situ monitoring, including a coaxial camera with 10000 Hz maximum frame rate for melt pool imaging. The corresponding wavelength of the camera is 850 nm.

The corresponding wavelength of the camera is 850 nm. Using custom-developed scan strategy design software, the digital commands that AMMT uses set precise laser beam position and laser beam power at 100 KHz, and synchronously triggers the MPM camera every 100 µs (or 10 KHz) [26].Two types of in-process data can be acquired during the build on AMMT: pre-loaded digital commands and in-situ monitoring data. Digital commands construct the complete laser beam control, including timestamp,341laser beam position, laser power, and sensor trigger. The in-situ data includes layerwise images (LWI) and melt pool monitoring (MPM) images. LWI is the optical image of the surface of each layer, which can be used as a reference for recoated powder layer surface quality, or solidified surface quality [27]. MPM captures the coaxial melt pool images during the build with a resolution of 8 µm/pixel. Figure 3 shows the conceptual model of the AMMT and the sample data.

MPM captures the coaxial melt pool images during the build with a resolution of 8 µm/pixel. Figure 3 shows the conceptual model of the AMMT and the sample data. Like the typical L-PBF process, the AMMT uses a galvo mirror scanner to steer and control the laser path. It also utilizes a custom-designed dynamic translating lens for laser flat-field correction, instead of an f-theta lens. This custom scanning lens is also optimized to enable diffraction-limited image quality for the MPM camera. The AMMT also employs a stationary LWI camera near the galvo mirror system, and flexible, multi-element illumination system. The LWI field of view covers the entire part surface, representing the final surface of each layer after powder recoating, and after the laser scan. MPM captures images of the incandescent light that radiates from the melt pool during the build. The white spot in the image center is the melt pool in the MPM example in Figure 3.

MPM captures images of the incandescent light that radiates from the melt pool during the build. The white spot in the image center is the melt pool in the MPM example in Figure 3. Previous research published the calibration results between the MPM image grayscale values to the equivalent radiant temperature at the measured wavelength of 850 nm [28].Figure 3. The AMMT uses different sensors to collect in-situ data: the coaxial MPM images, LWI, and scan commands.For this study, an experiment is designed to create two types of overhanging features in one AM part. This experiment creates four nominally identical parts within the same build on a wrought nickel alloy 625 (IN625) substrate cut to 100 mm x 100 mm x 12.5 mm. All four parts have the same geometry: a bounding box 5 mm x 9 mm x 5 mm, a 45o overhang feature and a cylindrical cavity. The powder material is mixed recycled and virgin IN625 powder. The build consists of 250 layers at 20 µm per layer.

The powder material is mixed recycled and virgin IN625 powder. The build consists of 250 layers at 20 µm per layer. The build employs a constant laser scan speed (800 mm/s) constant laser power (195 W). The scan strategy employs a stripe scan pattern with skywriting, and no pre- or post-contour scans. The general scan direction is designed to rotate 90° every layer. Detailed experiment descriptions can be found in [28]. Table 1 lists the properties of scan pattern and the abbreviation used in this paper.342Table 1. Description of the scan patterns used in this work. A ‘vertical’ or ‘horizontal’ scan track means the track is oriented in the Y or X direction, respectively.

Description of the scan patterns used in this work. A ‘vertical’ or ‘horizontal’ scan track means the track is oriented in the Y or X direction, respectively. And, ‘from right to left’ or ‘from top to bottom’ means the tracks area scanned sequentially in the X or Y direction.Scan PatternNameLayer 1, 5, 9, … Layer 2, 6, 10,… Layer 3, 7, 11, … Layer 4, 8, 12, ….Vertical scan track from right to left Horizontal scan track from top to bottom Vertical scan track from left to right Horizontal scan track from bottom to topV-R2L H-T2B V-L2R H-B2TFigure 4 shows the designed part dimensions and the key geometric features. This paper focuses on the overhang surface on top of the cylindrical hole feature with 3 mm depth. Figure 4(b) is the cross-sectional view of the part in YZ plane at a point along the axis of the cylindrical feature. As shown, the overhang feature is introduced at Layer 126 and gradually developed until Layer 226.

As shown, the overhang feature is introduced at Layer 126 and gradually developed until Layer 226. The red line marks the pure overhang surface on Layer 226. The solid and dashed lines indicate the infilling (laser on) and non-infilling (laser off) regions, respectively.Figure 4. The part dimensions and key geometric features. (a) the 3D model showing a horizontal cylindrical feature on one side of the part with 3 mm depth (and 4 mm diameter). (b) is cross-sectional plane showing relative layer positions. Solid lines represent the infilling region (laser on). Dashed line is the non-infilling region (laser off). The shaded region indicates un-melted powders within the hole. The initial overhang surface is red line on Layer 226.The following sections explain how MPM data are used to create remelting maps for overhang build process study.

The initial overhang surface is red line on Layer 226.The following sections explain how MPM data are used to create remelting maps for overhang build process study. For demonstrative purposes, this study only studies the overhang associated with the cylinder hole using the data from one part.MethodIn this section, we will introduce the method of creating a layerwise remelting map using in-situ MPM data. Figure 5 shows the general workflow of this procedure. It starts with data preprocessing to clean the raw data. Data registration is the second step to synchronize and align the data from various sources. Data fusion is the next step. The objective is to integrate the multimodal data into one indicator by applying various rules that depend on the purposes of the fusion, and/or the physical meanings of the data. After that, the fused AM data could be used for mapping the remelting conditions of each layer.

After that, the fused AM data could be used for mapping the remelting conditions of each layer. The final step is to calculate the remelting ratio.343The first step is to preprocess the raw data, which includes correcting the signal delay in digital command and removing noise from the raw MPM images. The signal delay is measured from sensor and machine calibrations [26]. Two features in the melt pool measurement are removed in data preprocessing: the background or dark-level noise, and the effect of plumes and spatters. A low threshold grayscale value, 20 in this work, is deployed to remove the background noise. The largest connected area (LCA) method is used to remove the spatters in the MPM images. Observations found that spatters typically have much smaller sizes than melt pools where they originate.

The largest connected area (LCA) method is used to remove the spatters in the MPM images. Observations found that spatters typically have much smaller sizes than melt pools where they originate. After the LCA filter, each melt pool image only contains one large white spot – the melt pool.Preprocess raw dataRegister in-situ dataApply AM data fusionCreate remelting mapCalculate remelting ratioFigure 5. Workflow for calculating the remelting ratio.Data registration is a concept in which data from various sources are synchronize or aligned within the same temporal or spatial domain. Spatial registration procedures transform different datasets to a single coordinate system [23]. The authors’ previous works developed an AM specific data registration method to register the pixels in each MPM image to a global coordinate system [24]. Detailed procedures could be found in Feng et al 2021 and Lu et al 2020 [24, 29].

Detailed procedures could be found in Feng et al 2021 and Lu et al 2020 [24, 29]. The registration maps the MPM image files to the commanded scanning position, which is defined in the AM machine and laser’s coordinate system. The registered MPM information still requires further processing conducted by users based on specific application needs. However, the MPM data fusion step plays a role in providing physical meanings to the MPM data by matching it to the spatial location within the part where and when each melt pool image was captured.Data fusion, as it stated, is “a technology to enable combining information from several sources in order to form a unified picture” [30]. It has been commonly used in defense but rare in manufacturing due to the lack of rich data types. New technological improvements continually increase the quantity of AM in-situ data acquired by built-in sensors and controllers [25].

New technological improvements continually increase the quantity of AM in-situ data acquired by built-in sensors and controllers [25]. Applying data fusion methods can provide advantages to this data-heavy AM domain. A multi-scale hierarchical AM data fusion method was recently introduced and demonstrated to address specific AM problems [11]. It fuses AM in-process raw data in pointwise, trackwise, layerwise, and partwise levels. Melting condition considered in this paper is mainly investigated with the spatial alignment at trackwise level. The next few paragraphs introduce the algorithm used to create remelting maps based on this AM data fusion strategy.Melt pool width and hatch spacing are two critical factors of remelting. Hatch spacing is a process parameter preconfigured during scan strategy development before the build starts. For a fixed hatch spacing, wider melting tracks have a higher chance to create more overlapping and remelting.

For a fixed hatch spacing, wider melting tracks have a higher chance to create more overlapping and remelting. Conversely, for a fixed melt pool width, wider hatch spacing decreases overlapping and remelting. The melt pool width can be directly measured after the build, which being verified in many single-track experiments [31]. However, it is not easy to measure the actual melt pool width from in-situ images in a multi-track layer. The actual melting and solidification344temperature may only be identified with a thermally calibrated imager [32, 33]. Figure 6 shows an example of melt pool width measurements based on different threshold values, and the imager pixel scaling of 8 μm/pixel. The left chart shows the width measured using different threshold grayscale values, and the right figure is an MPM image with a melt pool moving vertically from top to bottom.

The left chart shows the width measured using different threshold grayscale values, and the right figure is an MPM image with a melt pool moving vertically from top to bottom. While increasing the threshold value, larger areas within the melt pool image are removed (or set to binary value 0), and remaining areas of binary value 1 decrease in size. In this example, the thresholded region corresponding to the melt pool ‘tail’ becomes less elongated after threshold value of 150. Results in the rest of this paper are based on grayscale threshold 80, which is consistent to the authors’ previous work [4, 6].Figure 6. Melt pool width measurement based on different threshold value (left). A melt pool moving vertically from top to bottom. The contour represents the melt pool outline based on different threshold grayscale value. The width represents the widest horizontal length of contour.

The contour represents the melt pool outline based on different threshold grayscale value. The width represents the widest horizontal length of contour. The intersection of two red dashed lines is the centroid of the area bounded by the 255 value isoline, which could be different to the laser spot center [34].The key to creating remelting map, in general, is to find the overlapping region between adjacent melting tracks. When a laser beam scans, it not only melts the powder but also remelts the previously solidified tracks. MPM images capture the transient melting status. Figure 7 shows the schematic drawing of the overlapping region between two adjacent tracks, which can be calculated by fusing the registered in-situ MPM data. As shown in the " represents the nth figure, continuously sampled MPM images are aligned to the scan commands to form tracks. 𝑓! frame of the MPM images at Track i. The track number is extracted from the MPM images registered to the scan path.

𝑓! frame of the MPM images at Track i. The track number is extracted from the MPM images registered to the scan path. In the figure, T1 and T2 are the first and second melt pool tracks, respectively. The shaded area in red color is the overlapping region in between these two tracks. Note, the thresholding selected in the preprocessing stage can affect the area of the remelting region. However, this paper does not discuss what is the ideal threshold for the most accurate remelting consideration. Instead, it focuses on comparing differences of remelting for overhang and non-overhang surfaces as the same threshold applied on the whole part.345Figure 7. Overlapping region derived from MPM data registration and fusion.After constructing the remelting map for the entire layer using MPM images, a remelting index is calculated. Equation (1) shows how the remelting ratio is defined.

Equation (1) shows how the remelting ratio is defined. Note, due to the limitation of measurement technique, this paper cannot calculate the remelting overlay area along depth direction. This value indicates the proportion at which all the tracks are overlapped. A higher value indicates that tracks have strongly connected but may also suggests overheating. If this value is too small, it means that the part may suffer from “lack-of- fusion” defects.𝜑 = #! #" where φ is the ratio of the overlapping area 𝐴$ (green area in Figure 7), and total covered area by all tracks 𝐴%. The ratio is dependent on the melt pool threshold selected during the preprocessing stage. A higher threshold value would lead to lager melt pool area, which cause increment in both 𝐴$ and 𝐴%. For fixed hatching distance with multiple tracks, higher threshold value would produce a lower φ. For demonstrative purpose, this paper selects a relatively higher threshold value, 80, for all analysis.

For demonstrative purpose, this paper selects a relatively higher threshold value, 80, for all analysis. The following paragraphs would not further discuss this topic due to the lack of information about MPM image grayscale to melt pool width calibration.ResultA total of four parts are built in the experiment [28]. The result presented in this paper is based on Part 1and the complete overhang layers.Figure 8 shows a magnification of the remelting map for a regular layer (Layer 2) without an overhang from the top left corner (Track 1 and 2) and middle right (Track 26 and 27). The laser beam moves from right to left to create Track 1 and Track 27, where Track 2 and Track 26 are from the opposite direction under Track 1. Fused the registered MPM images create the outline of the melting tracks. The green shaded area is the overlapping region between the two tracks, namely the remelting of Track 1 and 2.

Fused the registered MPM images create the outline of the melting tracks. The green shaded area is the overlapping region between the two tracks, namely the remelting of Track 1 and 2. The middle figure plots the remelting map for the entire layer and the colormap marks the map from no melting (white) to 3 times remelting. 0 indicates the powder has been melted once but without further remelting. As shown, the majority of the map has a maximum one-time remelting. A few small spots are the area remelted twice caused by several oversized melt pools.346(1)Figure 8. The remelting map between Track 1 and Track 2 (top), and Track 26 and Track 27 (bottom). The remelting map of the entire layer (middle).Table 2 selects the same region on different layers for comparison, where the scan pattern, size, and location are the same. Layer 226 is the only layer in part that has a complete overhanging region (0.55	mm	 × 	3	mm). The cropped portion is built upon 200 layers of powders.

Layer 226 is the only layer in part that has a complete overhanging region (0.55	mm	 × 	3	mm). The cropped portion is built upon 200 layers of powders. Layer 10 is a regular layer without overhanging feature as a reference. Layer 250 is the last layer of the part. They use the colormap as Figure 8, where the white area indicates the powders not being melted. The last column lists the registered LWI at the same region of each layer. There is no significant difference observed between Layer 10 and 250. However, the remelting in Layer 226 is much lower, and the corresponding LWI seems very blurry without noticeable horizontal stripes.The remelting maps visually compare the melting conditions within a region which transitions from overhang to subsequent regular layers. Figure 9 aims to describe the impact from overhang features statistically by eight layers from the first overhang layer. The results are based on the same region of each layer as shown in Table 2.

Figure 9 aims to describe the impact from overhang features statistically by eight layers from the first overhang layer. The results are based on the same region of each layer as shown in Table 2. The red dots mark the average remelting index calculated from each layer corresponding the overhang region, and subsequent layers above this region. The open circles show the average remelting index of the entire infilling region from Layer 1 to Layer 25, with error bars indicating ±1 standard deviation. Those 25 layers has rectangular outline without overhang introduced. The lower and upper bounds are also plotted. As shown, the number decreased significantly for the immediate overhanging layer from 27.5% in layer 227, to 8.4% in layer347226. The actual overhang surface is geometrically eliminated from Layer 227. However, the same region remains under remelting. It was built upon post overhang surfaces.

The actual overhang surface is geometrically eliminated from Layer 227. However, the same region remains under remelting. It was built upon post overhang surfaces. The light blue bars plot the absolute difference between the overhang region to non-overhang region on current layer. The difference of 𝜑 between overhang region and non-overhang region on Layer 226 is 15.2%. Based on the data, it takes three layers to recover the remelting ratio back to the normal range. Table 2. Remelting map and corresponding LWI for the same region on different layers.Layer num 10Scan direction H-T2BDescriptionOne regular layerRemelting map for overhang regionLWI at the same region226250H-T2BH-T2BFirst overhang layer 25th regular layer above overhangFigure 9. The 𝜑 from Layer 226 to Layer 233 of the overhang region cropped from each layer. Red dots mark the 𝜑 in current layer. Error bar represents the average remelting range for regular layers with the same scan pattern.

Red dots mark the 𝜑 in current layer. Error bar represents the average remelting range for regular layers with the same scan pattern. The purple curve is the fitted average remelting ratio for Layer 226 to 233, which shows the recovering trend after overhang. The light blue bars are the absolute differences between overhang region and non-overhang region on current layer.Figure 10 compares an X-ray Computer Tomography (XCT) slice from volume data and remelting map of Layer 226. For clarification, the XCT slice is the preliminary result that taken approximately around 4.500 mm to 4.540 mm in Z direction, which roughly corresponds to layer 226. As shown, the dashed line circles the same region on this layer. The remelting map clearly shows the low overlapping between tracks. Similarly, the XCT scan shows this region has different grayscale than the area without overhang.

The remelting map clearly shows the low overlapping between tracks. Similarly, the XCT scan shows this region has different grayscale than the area without overhang. Note, the darkness of radiographic348density in the X-Ray image is usually affected by sample transmitted density. It is reasonable to assume this region has relative lower density than normal value.Figure 10. Comparison between XCT scan and remelting map of Layer 226. (a) is the XCT slice approximately between 4.50 mm to 4.54 mm in Z direction. (b) is the complete remelting map of Layer 226.Conclusion and DiscussionThis paper develops a method to calculate the remelting map based on AM data fusion techniques. The parts with overhanging surfaces were built to verify the method. The result shows the remelting ratio of overhang surfaces is significantly lower than the average of regular layers. An important finding is the irregular remelting conditions would not immediately recover in post-overhanging surface.

An important finding is the irregular remelting conditions would not immediately recover in post-overhanging surface. The following layers remain a lower remelting ratio. To receive full support, the part takes 3 layers to build the strength. The remelting ratio above the overhang region increases from layer to layer to approach that of the non-overhang regions. The proposed method has limitation in precise validation and remelting between layers.Future research will continue to investigate the physical changes caused by overhanging features, where this paper presents an expected result for complete overhang surfaces. It could be valuable to analyze the conditions for those “mild” overhanging surfaces. Data fusion can expand to multi-modal (MPM and LWI) and deep correlation between remelting ratio and part quality (porosity and surface roughness). Future research would also investigate the most appropriate threshold value to find the melt pool outline.

Future research would also investigate the most appropriate threshold value to find the melt pool outline. This step may potentially improve the accuracy of remelting ratio. One ongoing research from the coauthor is the recovering scan strategy. This strategy aims to use a near real-time control method to adjust such scan strategy as laser power and spot size in the next layer to fix the remelting issues on current layer.DisclaimerCertain commercial systems are identified in this paper. Such identification does not imply recommendation or endorsement by NIST; nor does it imply that the products identified are necessarily the best available for the purpose. Further, any opinions, findings, conclusions, or recommendations expressed in this material are those of the authors and do not necessarily reflect the views of NIST or any other supporting U.S.

Further, any opinions, findings, conclusions, or recommendations expressed in this material are those of the authors and do not necessarily reflect the views of NIST or any other supporting U.S. government or corporate organizations.AcknowledgmentsThis material is based upon work supported by the National Institute of Standards and Technology (NIST)under Cooperative Agreement number NIST 70NANB18H258.349References[1] Babuska, T. F., Krick, B. A., Susan, D. F., 2021, "Comparison of Powder Bed Fusion and Directed Energy Deposition for Tailoring Mechanical Properties of Traditionally Brittle Alloys," Manufacturing Letters, 28pp. 30-34. [2] Lopez, F., Witherell, P., and Lane, B., 2016, "Identifying Uncertainty in Laser Powder Bed Fusion Additive Manufacturing Models," Journal of Mechanical Design, 138(11) . [3] Arısoy, Y. M., Criales, L.

[3] Arısoy, Y. M., Criales, L. E., Özel, T., 2017, "Influence of Scan Strategy and Process Parameters on Microstructure and its Optimization in Additively Manufactured Nickel Alloy 625 Via Laser Powder Bed Fusion," The International Journal of Advanced Manufacturing Technology, 90(5) pp. 1393-1417. [4] Yang, Z., Lu, Y., Yeung, H., 2020, "From Scan Strategy to Melt Pool Prediction: A Neighboring-Effect Modeling Method," Journal of Computing and Information Science in Engineering, 20(5) . [5] Chen, H., Gu, D., Xiong, J., 2017, "Improving Additive Manufacturing Processability of Hard-to-Process Overhanging Structure by Selective Laser Melting," Journal of Materials Processing Technology, 250pp. 99- 108. [6] Yeung, H., Yang, Z., and Yan, L., 2020, "A Meltpool Prediction Based Scan Strategy for Powder Bed Fusion Additive Manufacturing," Additive Manufacturing, 35pp. 101383.

99- 108. [6] Yeung, H., Yang, Z., and Yan, L., 2020, "A Meltpool Prediction Based Scan Strategy for Powder Bed Fusion Additive Manufacturing," Additive Manufacturing, 35pp. 101383. [7] Yeung, H., and Lane, B., 2020, "A Residual Heat Compensation Based Scan Strategy for Powder Bed Fusion Additive Manufacturing," Manufacturing Letters, 25pp. 56-59. [8] Yeung, H., Lane, B., and Fox, J., 2019, "Part Geometry and Conduction-Based Laser Power Control for Powder Bed Fusion Additive Manufacturing," Additive Manufacturing, 30pp. 100844. [9] Patel, S., and Vlasea, M., 2020, "Melting Modes in Laser Powder Bed Fusion," Materialia, 9pp. 100591. [10] Sun, S., Brandt, M., and Easton, M., 2017, "Powder Bed Fusion Processes: An Overview," Laser Additive Manufacturing, pp. 55-77.

100591. [10] Sun, S., Brandt, M., and Easton, M., 2017, "Powder Bed Fusion Processes: An Overview," Laser Additive Manufacturing, pp. 55-77. [11] Yang, Z., Lu, Y., Li, S., Li, "IN-PROCESS DATA FUSION FOR PROCESS MONITORING AND CONTROL OF METAL ADDITIVE MANUFACTURING," International Design Engineering Technical Conferences and Computers and Information in Engineering Conference, Anonymous ASME, . [12] Tang, M., Pistorius, P. C., and Beuth, J. L., 2017, "Prediction of Lack-of-Fusion Porosity for Powder Bed Fusion," Additive Manufacturing, 14pp. 39-48. [13] Tenbrock, C., Fischer, F. G., Wissenbach, K., 2020, "Influence of Keyhole and Conduction Mode Melting for Top-Hat Shaped Beam Profiles in Laser Powder Bed Fusion," Journal of Materials Processing Technology, 278pp. 116514. [14] Kok, Y., Tan, X. P., Wang, P., 2018, "Anisotropy and Heterogeneity of Microstructure and Mechanical Properties in Metal Additive Manufacturing: A Critical Review," Materials & Design, 139pp. 565-586.

P., Wang, P., 2018, "Anisotropy and Heterogeneity of Microstructure and Mechanical Properties in Metal Additive Manufacturing: A Critical Review," Materials & Design, 139pp. 565-586. [15] Yang, Z., Lu, Y., Yeung, H., 2020, "3D Build Melt Pool Predictive Modeling for Powder Bed Fusion Additive Manufacturing," International Design Engineering Technical Conferences and Computers and Information in Engineering Conference, Anonymous American Society of Mechanical Engineers, 83983, pp. V009T09A046. [16] Yadroitsev, I., and Smurov, I., 2010, "Selective Laser Melting Technology: From the Single Laser Melted Track Stability to 3D Parts of Complex Shape," Physics Procedia, 5pp. 551-560. [17] Fox, J. C., Moylan, S. P., and Lane, B. M., 2016, "Effect of Process Parameters on the Surface Roughness of Overhanging Structures in Laser Powder Bed Fusion Additive Manufacturing," Procedia Cirp, 45pp. 131- 134.

P., and Lane, B. M., 2016, "Effect of Process Parameters on the Surface Roughness of Overhanging Structures in Laser Powder Bed Fusion Additive Manufacturing," Procedia Cirp, 45pp. 131- 134. [18] Wang, D., Yang, Y., Yi, Z., 2013, "Research on the Fabricating Quality Optimization of the Overhanging Surface in SLM Process," The International Journal of Advanced Manufacturing Technology, 65(9-12) pp. 1471-1484. [19] Calignano, F., 2014, "Design Optimization of Supports for Overhanging Structures in Aluminum and Titanium Alloys by Selective Laser Melting," Materials & Design, 64pp. 203-213.[20] Cheng, B., and Chou, K., 2015, "Geometric Consideration of Support Structures in Part Overhang Fabrications by Electron Beam Additive Manufacturing," Computer-Aided Design, 69pp. 102-111.350[21] Yang, Y. P., Jamshidinia, M., Boulware, P., 2018, "Prediction of Microstructure, Residual Stress, and Deformation in Laser Powder Bed Fusion Process," Computational Mechanics, 61(5) pp. 599-615.

P., Jamshidinia, M., Boulware, P., 2018, "Prediction of Microstructure, Residual Stress, and Deformation in Laser Powder Bed Fusion Process," Computational Mechanics, 61(5) pp. 599-615. [22] Mani, M., Feng, S., Brandon, L., 2017, "Measurement science needs for real-time control of additive manufacturing powder-bed fusion processes," CRC Press, . [23] Feng, S. C., Lu, Y., Jones, A. T., 2021, "Additive Manufacturing in-Situ and Ex-Situ Data Registration and Metadata Definition," Journal of Computing and Information Science in Engineering, . [24] Lu, Y., Yang, Z., Kim, J., "Camera-Based Coaxial Melt Pool Monitoring Data Registration for Laser Powder Bed Fusion Additive Manufacturing," . [25] Lane, B., Mekhontsev, S., Grantham, S., 2016, "Design, developments, and results from the NIST additive manufacturing metrology testbed (AMMT)," Solid Freeform Fabrication (SFF) Symposium, Anonymous pp. 1145-1160. [26] Yeung, H., Lane, B. M., Donmez, M.

1145-1160. [26] Yeung, H., Lane, B. M., Donmez, M. A., 2018, "Implementation of Advanced Laser Control Strategies for Powder Bed Fusion Systems," Procedia Manufacturing, 26pp. 871-879. [27] Scime, L., Siddel, D., Baird, S., 2020, "Layer-Wise Anomaly Detection and Classification for Powder Bed Additive Manufacturing Processes: A Machine-Agnostic Algorithm for Real-Time Pixel-Wise Semantic Segmentation," Additive Manufacturing, 36pp. 101453. [28] Lane, B., and Yeung, H., 2020, "Process Monitoring Dataset from the Additive Manufacturing Metrology Testbed (AMMT): Overhang Part X4," Journal of Research of the National Institute of Standards and Technology, 125pp. 1-18. [29] Feng, S. C., Lu, Y., and Jones, A.

1-18. [29] Feng, S. C., Lu, Y., and Jones, A. T., 2020, "Measured Data Alignments for Monitoring Metal Additive Manufacturing Processes Using Laser Powder Bed Fusion Methods," International Design Engineering Technical Conferences and Computers and Information in Engineering Conference, Anonymous American Society of Mechanical Engineers, 83983, pp. V009T09A022. [30] Hall, D. L., and Llinas, J., 1997, "An Introduction to Multisensor Data Fusion," Proceedings of the IEEE, 85(1) pp. 6-23. [31] Lu, Y., Yang, Z., Eddy, D., "Self-improving additive manufacturing knowledge management," Asme 2018 international design engineering technical conferences and computers and information in engineering conference, Anonymous American Society of Mechanical Engineers Digital Collection, . [32] Lane, B., Jacquemetton, L., Piltch, M., 2020, "Thermal Calibration of Commercial Melt Pool Monitoring Sensors on a Laser Powder Bed Fusion System,".

[32] Lane, B., Jacquemetton, L., Piltch, M., 2020, "Thermal Calibration of Commercial Melt Pool Monitoring Sensors on a Laser Powder Bed Fusion System," . [33] Lane, B., and Yeung, H., 2019, "Process Monitoring Dataset from the Additive Manufacturing Metrology Testbed (AMMT):" Three-Dimensional Scan Strategies"," Journal of Research of the National Institute of Standards and Technology, 124pp. 1-14. [34] Zhirnov, I., Mekhontsev, S., Lane, B., 2020, "Accurate Determination of Laser Spot Position during Laser Powder Bed Fusion Process Thermography," Manufacturing Letters, 23pp. 49-52.351.

alyzing the Cutting Process of a Heated Flexible Blade in Extruded Polystyrene FoamBram de Smit and Han J. Broek Faculty of Industrial Design Engineering, Delft University of Technology, Landbergstraat 15, NL-2628CE Delft, the Netherlands a.desmit@io.tudelft.nlReviewed, accepted August 30, 2004AbstractFor the “Free Form Thick Layered Object Manufacturing” technology that is being developed at the Delft University, it is vital that Extruded Polystyrene Foam (XPS) can be cut accurately with a heated blade. The shape of the blade is actively controlled during the cutting process, which results in double curved cutting surfaces. In order to make this cutting process controllable it must be known how the cutting behavior is under varying conditions, like cutting speed, cutting angle and heating power.

In order to make this cutting process controllable it must be known how the cutting behavior is under varying conditions, like cutting speed, cutting angle and heating power. The authors executed a range of experiments, analyzed the results and describe a practical model for the cutting process of XPS with a heated blade.IntroductionThe FF-TLOM technology is being developed for the use in an early stage of the design process for large objects [3]. It combines the advantages of layer oriented prototyping technologies, like a high form freedom and ease of process planning, with the advantages of decremental technologies like low material cost and high possible production speed. The specific properties of the process make it especially applicable in the domain of large freeform shapes [2], [4].

The specific properties of the process make it especially applicable in the domain of large freeform shapes [2], [4]. The models produced by the system can be used as concept models during the early design stage, however the possibility of using the model as a basis for a high quality presentation model, for example by covering it with modeling clay, can be considered. The FF-TLOM technology is based on the use of layers of extruded polystyrene foam. The layers are cut in a freeform than stacked. This stacking produces a 3 shape and dimensional physical model (figure 1). The technology can be divided in two research fields. The first field is about the decomposition of a freeform CAD model into producible layers [5]. The second field covers the production process, necessary to create the layers. The cutting tool is a part of the hardware needed for the production of layers.

The second field covers the production process, necessary to create the layers. The cutting tool is a part of the hardware needed for the production of layers. It exists of a flexible knife (blade) that can be heated, four actuators that can give the blade a curved shape, and a frame that connects the earlier mentioned parts. The process of cutting XPS with the mentioned flexible blade under varying conditions is being investigated this paper.Figure 1: Free Form model out of thick layers.591Earlier work on the topicFigure 2: Cutting experimentsNot much research is known to the authors, on hot blade cutting of polystyrene foam. Viswanathan et. al. studied the process of hot wire cutting with constant speed and power [7], More recently Ahn et. al studied the optimization of the hot wire cutting process [1].

Viswanathan et. al. studied the process of hot wire cutting with constant speed and power [7], More recently Ahn et. al studied the optimization of the hot wire cutting process [1]. During early explorations and brainstorm sessions we generated a list of ten parameters that are influential in the cutting process namely 1) the used foam material, 2) the blade temperature, 3) dissipated thermal energy, 4) blade cross-section, 5) cutting speed, 6) cutting force, 7) blade material, 8) bladetemperature distribution, 9) blade radiation pattern, and 10) maneuverability of the blade. Reasoning about these parameters and their influence we decided that parameter 1 would be kept constant and parameters 2, 3, 5, 6 and 8 would be experimentally explored [6]. The remaining parameters would be considered in a later stage when an optimal tool design is needed. An experimental set-up was created consisting of a straight blade that can be heated using a controllable amount of electric energy.

An experimental set-up was created consisting of a straight blade that can be heated using a controllable amount of electric energy. We first measured the thermal behavior of a blade when energy was supplied and made response charts concerning temperature changes, resistance changes and temperature distribution over the blade. Actual cutting experiments were executed in order to determine optimal cutting conditions for a specific type of extruded polystyrene foam (figure 2). The outcome of the experiments is a set of parameter combinations, each with an output value representing whether the resulting surface is bad, moderate or good. By combining the results in one graph we found a window within which the cutting process produces good surface results (figure 3). We see that the amount of thermal energy that is being dissipated by the blade must be roughly between 0.03 and 0.07 Joule per square millimeter foam surface.

We see that the amount of thermal energy that is being dissipated by the blade must be roughly between 0.03 and 0.07 Joule per square millimeter foam surface. This means that by a given output power the cutting speed may vary between Vmin and (7/3) * Vmin . The cutting forces in this parameter window are very small and in fact not measurable in our setup.Process Window for XPS cutting2y t i l a u Q e c a f r u S1000.10.120.110.080.09 Joule / mm27 0.00.040.050.130.030.020.010.060.140.150.16Figure 3: Cutting window. Quality level 0 is bad, 1 is moderate and 2 is good.Problem descriptionFrom the described experimental results we found a process window in which ‘good’ surface quality can be produced, with one type of blade, cutting in a tangential direction. This is however not enough information for a proper control of the cutting process under all circumstances (figure 3).

This is however not enough information for a proper control of the cutting process under all circumstances (figure 3). Apart from the surface quality it is necessary to be able to predict the exact position where the surface will be formed as a function of different process parameters. Therefore we must determine the offset of the produced surface from the centreSmooth surfaceRough surface due to altering cutting direction592Figure 4: Cutting direction influence0.17line of the blade, under varying conditions. The first step in this is to find a formula that can predict the width of the produced gap during the cutting process. Second we must be able to control the surface quality under varying conditions. Apart from the thermal power and the cutting speed there is a need for more insight in the effects of the blade geometry and the cutting angle. In the earlier experiments we did not study the offset of the surface from the centre line of the blade.

In the earlier experiments we did not study the offset of the surface from the centre line of the blade. Further we also did not go into the thermal behavior of the blade during the cutting process.Blade Geometry Influence In this text we distinguish the blade geometry and the blade shape. We define the blade geometry as the shape of the cross section of the material because the blade is an extruded cross section. The blade geometry can for example be rectangular or circular. The blade shape is the global shape of the blade, due to the parameters that affect the blade. The blade shape can be straight or curved with non-, one- or more inflection points. In our earlier experiment we used a cutting blade that was specially made for foam cutting with a handheld tool. It is a Nickel-Chromium steel with a specific cross section shape which includes a cutting edge. (Figure 5) This material is normally used to create curved cutting blades with a fixed shape.

It is a Nickel-Chromium steel with a specific cross section shape which includes a cutting edge. (Figure 5) This material is normally used to create curved cutting blades with a fixed shape. It is not very well applicable for our purpose because we want to continuously change the blade curvature. One of the things that we learned from the first experiments was that optimal cutting happens virtually foam material withdraws from the blade just before it is touched. This means that the stiffness of the blade, which was quite high for the original material, can be lowered considerably. The cross sections that we are going to examine will be rectangular without cutting edge because this is the best shape to make controlled bending of the blade possible. The width and height of the material will be varied in order to find the process influence of these values. The criteria for this optimal ratio are shape controllability and cutting performance.

The width and height of the material will be varied in order to find the process influence of these values. The criteria for this optimal ratio are shape controllability and cutting performance. The Geometry influence is strongly related to the Cutting angle. Therefore a sequence of measurements must be designed to determine the influence of the separate parameters.Perpendicular directionBladeTangential directionThicknessWidthCutting directionforceless astheFigure 5. The geometry (cross section) of the original cutting bladeAxial directionCutting directionPerpendicular directionCutting directionτTangential directionγTangential directionBlade cross sectionBlade top view593Figure 6: components of cutting directionCutting Angle The FF-TLOM principle makes it impossible to create the needed surfaces by straight cuts only.

When curved cuts have to be made and dynamically changing blade curvature is applied it is unavoidable that parts of the blade will cut through the material in a non-tangential direction. The cutting direction will usually have a tangential component, a perpendicular component and an axial component. (Figure 6) A number of test cuts must be performed to determine the effect of these cutting-direction influences. From the results of these measurements we may be able to describe a virtual tool shape. This virtual tool should describe the area around the blade where foam is removed, using a certain thermal energy, speed and direction.Surface offset The surface that is produced by a cutting action usually has not touched the cutting blade. The blade has actually passed the surface at a distance. This distance, the surface offset, is expected to be dependent on the momentary process parameters.

The blade has actually passed the surface at a distance. This distance, the surface offset, is expected to be dependent on the momentary process parameters. For a proper control of the produced surfaces we have to determine a model of the surface offset as a function of the relevant process parameters. Our experiments must be designed in such a way that we can derive such a model. The surface offset behavior is an important factor in the virtual tool description. Our measuring approach starts with the assumption that the offset is symmetrical on both sides of the blade. In this case we can determine the offset by making two test-cuts with exactly the same process parameters, parallel to each other at a well know distance. The resulting piece of material must have a remaining thickness of the nominal thickness minus two times the offset distance.

The resulting piece of material must have a remaining thickness of the nominal thickness minus two times the offset distance. (Figure 7) This means that by performing the needed test cuts two times in stead of one, we will automatically have the necessary information on the offset behavior.Blade temperature during cutting Until this moment we do not know what the blade temperature is during cutting. We have seen in our first experiments cutting that behavior changes during the cut. The resulting surface is different at the beginning of the cut than on the end of the cut. A reason for that can be a change in temperature during cutting, for example because some material does touch the blade and drains some thermal energy from it. To make this process better understandable and controllable we must measure the temperature of the blade during the cut. The temperature change may not be homogeneous along the blade, but can be different for the center of the foam and the edges of the foam.

The temperature change may not be homogeneous along the blade, but can be different for the center of the foam and the edges of the foam. Therefore we will measure the temperature on several points, distributed over the blade. From the knowledge of the temperature behavior we may be able to conclude whether and how much extra energy must be added to the blade during a long cut.Bottom offsetRemaining thicknessNominal ThicknessTop offsettheFoam materialFigure 7: offset measurementHypothesisFrom our earlier experiments and from others research [1],[6], we expect the gap width that occurs when XPS is being cut with a heated blade depends on: 1) the used foam material, 2) the594blade geometry, 3) the cutting angle γ, 4) the cutting angle τ, 5) the dissipated thermal power in Watt per mm. blade length P, and 6) the cutting speed in mm/sec V.

blade length P, and 6) the cutting speed in mm/sec V. In our earlier experiments a process window was found for the production of an acceptable surface quality, limited to one foam material and one blade geometry and with both the angles put to zero The surface offset was only explored roughly. The goal of this research is to find an offset predicting formula for the cutting process. It is expected that this will make a better-defined process window necessary. However the main target is the offset control. We will try to create offset formulas for one type of foam material and a limited number of geometries. This means that the surface offset σ will be described as:σ geom.A = F (γ,τ, P, V )P/V influence In our earlier experiments we combined the P and V parameter into one parameter Q, which represents the amount of thermal energy dissipated per square surface unit (Joule/mm2).

Evaluation of the surface quality in relation to this energy gave us a process window within which good surface quality was produced. At the start of our current experiments we will use the P and V parameters in combination; P/V. Our first hypothesis is that the offset is proportional with this value within the process window, which our area of interest. That means that the formula will contain a factor Q (= P/V).Influence of angle τ A change in the angle τ can be described as a change in the amount of dissipated thermal energy per surface unit. The change factor can be calculated from the change of the cutting length due to angle change. In figure 8 we see that the cutting part of the blade has a length a, witch is equal to the material thickness. If the blade is rotated over an angle τ the length of the cutting part of the blade is b, which can be calculated as b = a / cos τ.

If the blade is rotated over an angle τ the length of the cutting part of the blade is b, which can be calculated as b = a / cos τ. This means that the formula for surface offset will contain a factor 1/cos τ.τbaFigure 8: Influence of angle τInfluence of angle γ We expect that a change of γ from zero degrees will in some way influence the surface offset. The amount of offset increase must be experimentally determined. We do not know whether the influence is proportional or otherwise. Because the amount of thermal energy per surface unit is not changed the influence may be small. It could even be the case that only a change in surface quality is found. In our hypothesis for a formula we will use a factor (γb+1), in which b is a constant that has to be determined from our experimental results.

In our hypothesis for a formula we will use a factor (γb+1), in which b is a constant that has to be determined from our experimental results. When we combine the expected influences into one formula we find:P V1 cosσa ⋅=⋅⋅τ595( γ)1 b ++⋅cIn this formula we expect that constants a and c are material related, while constant b should be blade geometry related. It is our aim to explore whether we can make this formula applicable for our process window.The experimental set upThe experiments that we performed include only straight cuts, but with different blade orientations, and with different speed settings. In our earlier experiments we used a three axis milling machine as the moving machine for the cuts. This setup had some drawbacks, the most important was the problem that the axes are driven by stepper motors with a relatively course resolution. In the case of low speeds this means that the motion profile is not smooth but stepped.

In the case of low speeds this means that the motion profile is not smooth but stepped. Further the length of the motions is limited, and in our earlier experiments we found that we would have to make longer cuts as many of the cutting processes in the first set of experiments seemed not to be stabilized at the end of the test piece. Thirdly we wanted to be able to orient the in a controlled manner. Therefore we blade implemented a set up with a servo driven linear axis that can travel 1000 mm. This axis can accurately be controlled in speed without stepper effect, and is used to move the material though the cutting tool. The material is fixed to the linear axis by a vacuum- gripping device. The cutting tool is a fixed blade holder that is attached to a six degrees of freedom robot arm. The robot is not moving during the cut, but is used between the cuts to position and orient the tool relative to the material (figure 9).

The robot is not moving during the cut, but is used between the cuts to position and orient the tool relative to the material (figure 9). During the experiments three geometries, described by width * thickness, of stainless steel blades have been used: The C-type blade was used for a full set of speed and power measurements with angle, variations. A second blade with the C geometry was fitted with 6 temperature sensors (figure 10) and a subset of the first full set of measurements was performed to study the thermal behavior. With the P and Q geometries subsets and variants of the first full set were carried out in order to study the influence of different geometries. The foam material that was applied is Dow Styrofoam LB with a measured density of approximately 36 kg/m3. The cutting length of the slabs was 800 mm. and the slabs where 50 mm. thick.

The foam material that was applied is Dow Styrofoam LB with a measured density of approximately 36 kg/m3. The cutting length of the slabs was 800 mm. and the slabs where 50 mm. thick. The slices that were cut have a nominal thickness of 30 mm.Figure 9: The used experimental set upFigure 10: A blade fitted with temperature sensorsBlade type C P QGeometry 4 * 0.3 mm. 2 * 0.3 mm 3 * 0.3 mm596ResultsWe produced an amount of about 100 test pieces, each under specified conditions. We used 3 blade geometries. One of the geometries was used in two sets of measurements, one without and one with added temperature sensors. By doing so we tried to have a good insight in the temperature behavior during the cutting process. From the produced pieces we could measure the produced gap width and observe the surface quality.Measurement accuracy In order to determine the effect of blade parameters on the produced gap width in the foam we measured the thickness of test parts.

These parts are produced in two cuts, on for the top surface and one for the bottom surface (Figure 7). The thickness measurements were carried out with a slide gauge. A subset of the thickness measurements was repeated a number of times in order to find a value for the repeatability of the measuring method. We measured nine foam pieces each five times. In this cycles we found for one measurement point a maximum difference of 0.1 mm. The maximum standard deviation that we found was 0.04. We state that the repeatability of the physical measurement is 0.1 mm.Experiment repeatability It can be expected that the experiments that were carried out can only be reproduced with a limited accuracy. In order to find this repeatability we performed two types of repeatability tests.

In order to find this repeatability we performed two types of repeatability tests. First we selected a number of experiments that were repeated in another measuring session, that means another day, equipment has been switched off and on again, another piece of foam material (We made sure that all the used foam material is from the same production batch). Second we took one specific measurement which was repeated a number of times during the same experiment session, so at the same day, in the same piece of material with all the equipment left in the same mode. In the first group of repeated experiments we found varying results. We can observerepeated tests65.55h t d w p a gi4.543.532.52C 002C 009C 018C 026P 001P 001aP 002P 002aP 003 P 003atest codeFigure 11: Tests repeated over different sessions.

The first character in the code represents the Blade geometry used, varying angles and speeds were tested.h t d w p a g d e r u s a e mi4.454.44.354.34.254.24.154.14.0543.95Q005Q005ablok repeatabilityQ005bQ005ctest codeQ005dQ005eQ005fstartcentreendin the graph in figure 11 that the largest deviation between two repeated tests is 0.35 mm. This is the case between test C018 and C026. These are tests with the blade material that was tested the most extensive, and the experiments were distributed over several log in days. Verification learned that experiment C026 was performed at the beginning of the last session with this material. We found that that it was noticed that the blade was visibly corroded, as a result of the earlier cutting experiments. This is the only case where this observation was made. If wethe experimentfigure 12: Repeatability in one session.

This is the only case where this observation was made. If wethe experimentfigure 12: Repeatability in one session. All tests used the Q geometry and constant speed and angle settings.597startcentreendassume that this may have influenced the experiment significantly and therefore leave this case out we are left with a surprisingly good repeatability. There is one case left where at the start of the cut the deviation is 0.2 mm., in all other cases it is better than 0.1 mm. which was found as the actual measurement accuracy In the second repeatability test (Figure 12) we repeated a test 7 times in a row, we found values that are in a comparable magnitude. The maximum difference found there is 0.25 mm. In this group, test Q005e has the highest deviation. If we leave this measurement out the maximum difference found is limited to 0.15 mm..

The maximum difference found there is 0.25 mm. In this group, test Q005e has the highest deviation. If we leave this measurement out the maximum difference found is limited to 0.15 mm.. We state that the repeatability of the experiments overall is better than 0.25 mm.Power distribution over the blade during the cutting process. During the experiments it proved to be difficult to control the cutting parameters in an accurate manner. Being a process that is determined by the added amount of thermal energy it is crucial to be able to control this amount as precise as possible. One of the problems is that the dissipated power may be influenced by the actual cutting process itself. If we observe a graph of the blade temperature during a cutting action we find that the blade temperature is decreasing considerably in the piece of the blade that is passing through the foam material. We know that the specific resistance of a metal depends on the temperature of that material.

We know that the specific resistance of a metal depends on the temperature of that material. During the experiments we observed that the current that is consumed by the blade typically increases during the cut. This is understandable if we consider that the resistance of the blade decreases when the material is cooled down, and that the current is generated by a poorly stabilized power source. In order to determine the influence of this effect we studied a specific case: During a cutting test T002 we logged the temperature of the blade and the consumed current through the blade. (Figure 13) The length of the blade is 175 mm. We find that the temperature of the centre of the blade starts at 580 degrees Celsius, before the cut, and drops to 440 degrees Celsius during the cut. A thermocouple that is mounted just outside the cutting part of the blade shows virtually no temperature drop.

A thermocouple that is mounted just outside the cutting part of the blade shows virtually no temperature drop. We logged the current and overall voltage during the cut and found that the resistance of the blade is158.8 ·10-3 Ohm before the cut and 155 ·10-3 Ohm during the cut. That means that the resistance decreased by 3.8 ·10-3 Ohm. As mentioned the total length of the blade is 175 mm., while the cutting part of the blade is limited to 50 mm. Because the temperature is virtually not influenced outside the cutting part, we can assume that the change in resistance takes place in the cutting part of the material. The resistance of that part was (50 * 158.8)/175 = 45.37 ·10-3 Ohm. The resistance during the cut is than 45.37 – 3.8 = 41.57 ·10-3 Ohm. With the measured current of 19.65 A before the cut and 20.0 A during the cut this results in a dissipated power in the cutting part of the blade of 17.52Watt before the cut and 16.63 Watt during the cut.

That means that the power dissipation in the cutting part of the foam has dropped 5% during the cut, caused by the cooling of the blade. This is a considerable amount that has to be taken into account when a dedicated control device for the cutting power is implemented. For this paper we assume that the effect is working in all the executed tests in a comparable manner. This means that all the measured power values should have an offset of about 5 %.

For this paper we assume that the effect is working in all the executed tests in a comparable manner. This means that all the measured power values should have an offset of about 5 %. For the principles of the process this doesn’t make a large difference.T00220.1700650T in Foam20T out foam600current(A)19.9s u i c l e C s e e r g e D550500)19.8A(t n e r r u c45019.740035019.630019.525020019.444:16.543:44.243:11.942:39.642:07.341:35.044:48.7timeFigure 13: Thermal behavior of the blade during a cutting experiment.598AnalysisInfluence of power per square unit We performed a number of tangential cutting experiments with different blade geometries and power/speed combinations. When the resulting gap measurements are combined we find a graph in which we clearly observe a linear relation between the amount of dissipated thermal energy per square surface unit and the produced gap width (figure 14).

Neither the geometry that is applied nor the actual blade temperature is of any influence on the gap. From the graph we can find that in our formula:.m m nih t d w654321000.010.020.03gap widht vs. Q0.040.050.060.070.08Geo-CGeo-PGeo-QQ in Joules / mm2σa ⋅=P V⋅1 cosτ⋅( γ)1 b ++⋅cFigure 14: The gap width as a function of dissipated power for different blade geometriesThe constant a is 74.843 and the constant c is –0.2868Varying Tau with and without speed compensationInfluence of the angle τ As explained before a change in the angle τ can be described as a change in the amount of power per square surface unit. We did two experiments to verify this behavior. First we varied the angle τ without any speed compensation. And measured the resulting graph. Figure 15 shows the graph.

We did two experiments to verify this behavior. First we varied the angle τ without any speed compensation. And measured the resulting graph. Figure 15 shows the graph. Then we applied comparable τ the speed by variations but compensated) .m m( h t d w p a gi76.565.554.543.5speed constantspeed compensated30102030405060701 τcosTau (degrees)multiplying it with the factor. ThisFigure 15: Influence of angle τ on the gap widthresulted in the second curve shown in figure 15. From the graphs we conclude that our hypothesis on the influence of angle τ is correct.Influence of the angle γ We performed two sequences of experiments with varying γ and constant power dissipation. The first sequence used blade geometry C and the second used blade geometry P. Figure 16 shows the graph of the resulting gap widths. It is observable that both the graphs are horizontal with small deviations until a certain point.

Figure 16 shows the graph of the resulting gap widths. It is observable that both the graphs are horizontal with small deviations until a certain point. For The C geometry this point comes sooner than for the P geometry. The reason is that at this point the cutting process is no longer stable. The blade is touching the foam and a force is building up. The blade is pushed away, the force increasesm m nit h d wi43.93.83.73.63.53.43.33.23.13010Gap with with varying gamma2030Gamma4050Geo-CGeo-P599Figure 16: Influence of Angle γ on the gap widthfurther and the cut is no longer controllable. The observation that the two curves in the graph are about 0.25 mm. apart is explained by a slight difference in applied power. If a correction for the power dissipation is made they fit well together. The conclusion from this graph is that the angle γ has no influence on the gap width, up to a point where the cutting process becomes unstable.

The conclusion from this graph is that the angle γ has no influence on the gap width, up to a point where the cutting process becomes unstable. This point is earlier reached with a wider blade (the C geometry). Therefore the factor b in the formula is found to be zero.Asymmetry in the gap width Although the gap width is not influenced by the angle gamma, there is another effect that should be mentioned. During the experiment we observed that the offset of the surface from the centre line of the blade is not symmetrically distributed when the blade is cutting with a gamma angle. Further we observed that the surface quality differs significantly depending on this gamma angle. In figure 17 we can see that with the given blade angle and cutting direction the offset above the blade (Top offset) is larger than below the blade (bottom offset).

In figure 17 we can see that with the given blade angle and cutting direction the offset above the blade (Top offset) is larger than below the blade (bottom offset). Further it is marked that the surface above the blade has a glazed quality, which is not optimal for our purpose, while the surface below the blade is a clean high quality surface. If we invert the cutting direction in this picture we find that the offsets and surface qualities also swap to their opposites. In order to find more about the offset asymmetry we performed some tests with varying cutting directions. We defined the direction that was used for the majority of the measurements as the Normal direction, while the opposite direction is called Reverse direction. With each blade geometry we cut four test pieces, all with a γ of 30 degrees, but with all possible combinations of normal and reverse cutting directions.

With each blade geometry we cut four test pieces, all with a γ of 30 degrees, but with all possible combinations of normal and reverse cutting directions. From the graph in figure 18 we can conclude that there are only small differences between the NN and RR combination, namely maximum 0.15 mm. We expect them to be the same because it is in fact a repeated test, which is only performed in opposite direction. De 0.15 deviation is within the repeatability margins. The RN and NR combination show large differences, which is again expected because one of the two combines two large offsets and the other combines two small offsets. We found for the C geometry a large offset of 116% of the average, for the Q geometry we found a large offset of 117% of the average offset and for the P geometry we found a large offset of 108% of the average. The used power dissipation is not the same for the different geometries.

The used power dissipation is not the same for the different geometries. From this can we can conclude that there is indeed a significant asymmetry in the offset when a γ angle is used. We cannot yet predict this offsets in an accurate manner. Some extra cutting tests must be carried out to collect the needed data for that.Glazed surfaceTop offsetCutting directionBottom offsetClean SurfaceBladeFoam materialFigure 17: Effects of a cut made with an angle γ of the bladegap asymmetry for Gamma is 30765h t d w p a g4Geo-CiGeo-P3Geo-Q210RRNRNNRNcutting directionsFigure 18: calculated gap width for different combinations of cutting directions600ConclusionsThe formula that we created in our hypothesis has been tested and found to be correct. For the constants a and c we found applicable values, the constant b proved to be zero.

For the constants a and c we found applicable values, the constant b proved to be zero. The resulting formula is therefore:σ=(74.843⋅P V⋅1 cosτ.0) −2868It was found that it is crucial to have a precise control on the dissipated power in the blade. This is difficult, as the cutting process itself is influencing the amount of dissipated energy in the blade. We observed that when an angle γ is applied, the gap that is created is not symmetrical, seen from the centre line of the blade. More experiments are needed to determine a formula that predicts this asymmetry. In our experiments we found differences from 108% to 117% between the average offset and the largest offset in an asymmetric case with angles γ f 30 degrees. Finally it was found that the asymmetric gaps that are created when an angle γ is applied also have different quality levels on the two opposite sides of the gap.

Finally it was found that the asymmetric gaps that are created when an angle γ is applied also have different quality levels on the two opposite sides of the gap. This is an important finding in relation to the cutting strategy that should be followed in order to create optimal models.References[1] Ahn, D.G.. Lee, S.H., and Yang, D.Y., Investigation into thermal characteristics of linear hotwired cutting system for variable lamination manufacturing (VLM) process by using expandable polystyrene foam, International Journal of Machine Tools & Manufacture 42 (2002) 427-439.[2] Hope, R.L., Jacobs, P.A., and Roth, R.N., (1997), Rapid Prototyping with SlopingSurfaces, Rapid Prototyping Journal 3-1, pp. 12-19[3] Horváth, I., Broek, J.J., Lennings, A.F., and Vergeest, J.S.M., Reshapeable Concept Models by FF-TLOM: A True Alternative of Conventional Clay Modeling, ISATA'99, paper 99ME069[4] Horváth, I., Vergeest, J.S.M., Rusák, Z.

ALYZING THE DMLS-PROCESS BY A MACROSCOPIC FE-MODELF. Niebling, A. Otto, M. GeigerChair of Manufacturing Technology, University Erlangen-Nuremberg, GermanyAbstractThe presented macroscopic FE-model allows to analyze the thermal fields and the resulting stress built up during Selective Laser Sintering. Process and material parameters are focused on Direct Metal Laser Sintering (DMLS). The FE-model is introduced and the assumptions for the model are given. Three different geometric models are discussed. The 3D- model shows the sintering of a single line, whereas 2D-models are used for longitudinal and crosscuts of the sintering process. Aim of the investigation is a more basic knowledge about the process, which will lead to a stabilization and optimization of the process.IntroductionSelective Laser Sintering (SLS) is an additive Rapid Prototyping technology [1]. A 3D- computer model is sliced into layers and the layer information is transferred to a SLS-machine.

A 3D- computer model is sliced into layers and the layer information is transferred to a SLS-machine. This machine transmits the adequate layer information onto a powder bed, whereby the topmost layer is locally sintered, coherent to the sintered layers beyond. By reiteratively adding new powder on top and locally sintering with the energy of a laser beam a 3D part can be built. This technology was originally developed for making plastic prototypes at the University of Texas, Austin in 1987 [2]. The SLS-technology is today used for making plastic, metal and ceramic prototypes.Metal SLS-prototypes can be produced by two different ways [3], the Indirect Metal Laser Sintering process (IMLS), developed by DTM and the Direct Metal Laser Sintering process (DMLS) of EOS [4, 5, 6].

Both processes use a binder and a structure component in the powders and the relatively to the structure component low melting binder component is molten by the energy of the laser beam.The ILMS process by DTM uses polymer binders, which are burned out in following ovenprocesses and the porous material is infiltrated with metal.The DMLS process of EOS uses a powder mixture, consisting of metals with different melting points. Several powder mixtures are available from EOS.

Several powder mixtures are available from EOS. The bronze based powders can be processes in air whereas the steel based powders need an inert gas atmosphere during processing.The presented FE-model is based on parameters of the DMLS process and for an easier experimental accessibility for the described investigations the bronze based powder Direct Metal 50-V2 [7] was chosen, but in principle it is possible to adapt the model to every variant of the SLS process.Modeling the sintering behaviorAs a simulation tool for the FE-model Sysweld [8] from ESI was used. It allows the implementation of so called phases with different temperature dependent material properties and a time and temperature dependent function can be implemented for a phase change. These features were used for the simulation of the SLS-process: One phase was assigned to the material384parameter of the powder and another phase was assigned to the material parameters of the sintered material.

Furthermore, a function was implemented for the transformation from powder to sintered material. It converts the powder phase at the melting point of the binder (850 °C for Direct Metal 50-V2) to sintered material. Material properties of both phases meet at the melting point of the binder phase and have at this temperature properties of the molten material.For the simulation of the thermal field, density and heat conductivity were measured at room temperature for powder and sintered material. Especially the heat conductivity depends strongly on the temperature and it differs by a factor of 60 for powder and sintered material at room temperature. For the powder, the laws derived by [9] were used to model the heat dependency. The temperature dependency of the sintered material was approximated by applying the same gradient as of a similar bronze.

The temperature dependency of the sintered material was approximated by applying the same gradient as of a similar bronze. The specific enthalpy was derived from the specific heat capacity and the latent heat of the melting reaction of the fractions of the powder, weighted by their ratio.The simulation of the thermomechanical reactions of the SLS-process requires in addition to the parameters above the temperature dependent material parameters for yield strength, elasticity modulus and expansion coefficient. These were measured for the sintered material. Yield strength for powder is set to 10 MPa, because a lower yield strength caused instabilities in the numerical calculations.Finally, the heat coupling between laser beam and powder bed were modeled. A surface distributed energy coupling is not appropriate for metallic powder, as described in [10], so an approach similar to [11] with a volumetric heat source was used, fig. 1.

A surface distributed energy coupling is not appropriate for metallic powder, as described in [10], so an approach similar to [11] with a volumetric heat source was used, fig. 1. The diameter of the heat source corresponds to the gaussian beam diameter at the interaction zone and was set to 400 µm. The depth of the heat interaction zone equals the average grain size and was set to 50 µm. The verification of the model is given in [12]. The approach is in good consistency with the experimental studies.Fig. 1: Model of the energy coupling of the laser beam into the powder bed385Layer processing via SLS and derived geometric models for the FE-simulationThe top layer with the thickness h is locally exposed during the SLS-process to the energy of the laser beam corresponding to the layer information from the sliced computer model. The principle of layer exposition can be divided into two processing steps, fig.

The principle of layer exposition can be divided into two processing steps, fig. 2.Figure 2: Principle of laser exposition in the SLS-processFirst, the contour is sintered with the energy of the laser beam which will form the side surface of the part. The process parameters of this line set important properties for the final part like e.g. the roughness. A main process parameter is the line energy EL given by:E =LP L vS,where PL is the laser power and vS is the scan speed. The line energy is responsible for the amount of locally molten material and sets the local cooling behavior as described in [12].Secondly, the inner regions of the part are sintered by successive scanning over the powder bed. Besides the line energy, the area energy EF has a main influence on the final part quality. The area energy is given by:EF=hSP L ⋅ vS.The area energy is the line energy divided by the offset between the successive scans, called the hatching distance hS.

The area energy is given by:EF=hSP L ⋅ vS.The area energy is the line energy divided by the offset between the successive scans, called the hatching distance hS. The influence of the area energy on the part quality can be deduced from investigations in [13].386The FE-simulation of the complete processing within a 3D-model doesn’t make sense at the state of the art of the accessible computer power and software solutions. Therefore the process was investigated with different geometric models: The processing of the contour is examined by a 3D-model, which models the sintering of a single line. Basic conclusions can be drawn from this model about the thermal field and the stress built up. The complex process of filling the part by iterative scanning on the inner regions can be investigated by 2D-models of a longitudinal cut and a crosscut through the processed part.

The complex process of filling the part by iterative scanning on the inner regions can be investigated by 2D-models of a longitudinal cut and a crosscut through the processed part. The mechanical model is simulated as plane strain to show up the theoretical maximum of stress generation within the processing.Thermal and stress field for sintering linesA 3D-model as shown in fig. 3 was used to gain knowledge about the processing of the contour line and for acquiring basic information about the thermal and stress field during the sintering process. On a sintered stress free part with 4 mm length and a height of 1 mm a powder layer of 0,1 mm thickness is applied. The part has ambient temperature in it’s initial condition. The sintering process starts by the laser beam activation and the laser is moved over the powder bed for a length of 3 mm. Because of symmetry conditions only one half of the part was modeled.

The sintering process starts by the laser beam activation and the laser is moved over the powder bed for a length of 3 mm. Because of symmetry conditions only one half of the part was modeled. 35552 nodes and 31500 elements were used in the spatial FE-model.Figure 3: Temperature distribution embedded in geometric model for sintering a lineFig. 4 shows the temperature and stress development for a node in the bottom of the powder layer on the symmetry plane, marked in fig. 3. In the region between the layers delamination as a process failure can occur. It is therefore especially interesting, to study the temperature and stress development in this region.For a time span of about 3 ms the temperature locally exceeds 850 °C, the melting temperature of the binder material. The binder is molten and in the FE-model a phase change387from powder material to sintered material takes place in this region.

The binder is molten and in the FE-model a phase change387from powder material to sintered material takes place in this region. Material properties are at these temperatures those of the molten material, visible e.g. in the decreased yield strength.1800801600 °C70 MPa140060120050e r u t a r e p m e T1000800600400Temperature σσσσ xx σσσσ yy σσσσ zz Von Mises Stress Yield Stress40302010s s e r t S20000100510152025ms 3035TimeFigure 4: Temperature and stress development between the layers in the middle of a sintered trace. Process parameters: PL: 151 W, vS: 294 mm/sIn the following cooling tensile stresses increase, mainly in the longitudinal direction to the sintered trace, caused by strain obstruction of the contracting sintered trace. The von Mises stress builds up and reaches asymptotically a constant value of about 60 MPa. The resulting internal stresses can cause warpage of the part, if heated successively.

The von Mises stress builds up and reaches asymptotically a constant value of about 60 MPa. The resulting internal stresses can cause warpage of the part, if heated successively. Furthermore, in the early local cooling stage, the von Mises stress gains similar values like the yield stress and the elastic deformation ability of the material about 5 ms after solidification is nearly depleted. The buckling in the yield stress curve at 400 °C is due to a change in the tangent of the yield limit curve at these temperatures. The material behavior at temperatures beyond 400 °C is brittle which means that internal tensile stresses with a von Mises stress loading near to the yield limit can cause a crack initiation. The results of the simulation aim therefore to a possible origin of the delamination in the early local cooling.Longitudinal cut through the building process: Influence of scan vector length on temperature developmentPart properties as e.g.

density depend on the scan vector length, [14]. The scan vector length influences therefore the local thermal field during the sintering process. The thermal field can be described locally by the maximum temperature and the successive cooling behavior.The longitudinalcut FE-model was used to study this effect. The geometry of the model isshown in fig. 5. It has 12768 nodes and 12540 elements. The model is described in detail in [12]. In the following only the result of the dependency of the maximum temperature on the scan vector length is shown.388Fig. 5: Geometric model for the longitudinal cut through the sintering process. Shown number of elements is reduced.The dependency of the maximum temperature on the scanning way was studied by a parameter variation. Scan speed and line energy were varied, fig. 6. The maximum temperature is attained at about the same scanned distance, independent of the scan speed and with a light dependency on the line energy.

Scan speed and line energy were varied, fig. 6. The maximum temperature is attained at about the same scanned distance, independent of the scan speed and with a light dependency on the line energy. In the given example, at about 0,4 mm, which corresponds to the diameter of the laser beam on the powder bed. The temperature field is stabilized in the molten area at the interaction zone at that point. Therefore, it can be assumed that the dependency of material properties on the scan vector length is mainly due to a different local cooling behavior, caused by the different shape of the sintered upper layers.

In the sintering process heat is dissipated mainly in the sintered body due to the strong differing heat conductivity of metal powder and sintered body.1800°C1600140012001000Scan speed: 500 mm/s; Standardized line energy: 1e r u t a r e p m e T800600400200Scan speed: 400 mm/s; Standardized line energy: 1Scan speed: 320 mm/s; Standardized line energy: 1Scan speed: 320 mm/s; Standardized line energy: 1,25Scan speed: 320 mm/s; Standardized line energy: 1.56000,20,40,60,811,21,41,6mm 1,82Scanned wayFig. 6: Maximum temperature vs. scanned distance. Scan speed and line energy varied.389Crosscut trough the building processGeometric measures of the crosscut model are shown in fig. 7. This model is used for studying the effects of scanning strategies on the thermal field during processing and the resulting stress field. The geometric model has 3752 nodes and 3630 elements.Fig.

This model is used for studying the effects of scanning strategies on the thermal field during processing and the resulting stress field. The geometric model has 3752 nodes and 3630 elements.Fig. 7: Geometric model for the crosscut through the sintering process. (Shown number of elements is reduced.)In fig. 8, the von Mises stress distribution of a cooled part, sintered on a platform as a result of different scanning strategies is shown. Both scanning strategies were simulated with the same resulting area energy EF. Within the left scanning strategy, the stress distribution in the part is textured and the high stress loaded regions are at the edges. Together with the notch effect of the side shape results a critical processing leading to a high possibility for delamination. The example on the right side displays the result with a changed scanning strategy. The stress field is more uniform and areas of highest internal stresses are not near to the edges anymore.Fig.

The example on the right side displays the result with a changed scanning strategy. The stress field is more uniform and areas of highest internal stresses are not near to the edges anymore.Fig. 8: Resulting von Mises stress distribution of a sintered part on a platform for different scanning strategies. EF was kept constant.390SummaryA macroscopic FE-model is presented, which is capable of analyzing the transient temperature and stress field of the DMLS process. The model describes the powder and the sintered body with temperature dependent material properties and the process of local powder melting and cooling in a sintered condition. Results of a 3D-model of sintering a line indicate, that a possible crack initiation, which results in the process failure delamination is initiated in the early local cooling stage.

Results of a 3D-model of sintering a line indicate, that a possible crack initiation, which results in the process failure delamination is initiated in the early local cooling stage. Results of the 2D longitudinal cut FE-model indicate, that geometric size dependent effects like different porosity for identical machining parameters are due to changes in the cooling behavior. A significant change of the maximum temperature can not be observed for a scan length more than 0,4 mm, which means a stabilization of the heat input and output. Within a 2D-cross section model the influence on the stress development of scanning strategies can be studied. The macroscopic FE-model therefore gives basic predictions for a stabilization and optimization of the DMLS process.AcknowledgementThese investigations were financially supported by the DFG.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperANALYZING THE TENSILE, COMPRESSIVE, AND FLEXURAL PROPERTIES OF 3D PRINTED ABS P430 PLASTIC BASED ON PRINTING ORIENTATION USING FUSED DEPOSITION MODELINGR. Hernandez, D. Slaughter, D. Whaley, J. Tate, and B. Asiabanpour Ingram School of Engineering, Texas State University, San Marcos, Texas 78666AbstractTo achieve the optimum functionality and mechanical properties in the AM-based parts, it is vital to fully characterize parts under static mechanical loadings (tension, compression, and flexure) that are built in different orientations.

This research reports the results of the compression (ASTM standard D695), 4-point flexure (ASTM D790), and tensile (ASTM D 638 Type I) tests on the ABS plastic specimens that are designed according to the ASTM standards and are built in different orientations using the uPrint SE Plus 3D printer. This study examined the effects that printing 3D parts in different orientations (build angles) has on the mechanical properties of ABS P430 plastic. A total of 45 samples (15 tension, 15 compression, and 15 flexure) were printed in 5 orientations; 0 degrees in the XY plane, 45 degrees in the XY plane, 90 degrees in the XY plane, 45 degrees in the Z plane, and 90 degrees in the Z plane. The hypothesis was that the samples printed 0 degrees in the XY plane would be the strongest in compression and flexure, and also have the greatest modulus of elasticity.

The hypothesis was that the samples printed 0 degrees in the XY plane would be the strongest in compression and flexure, and also have the greatest modulus of elasticity. The samples printed 90 degrees in the XY plane were predicted to be the strongest in tension, having the largest tensile strength and lowest modulus of elasticity. The findings showed that printing 90 degrees in the XY plane resulted in the highest tensile strength compared to the other orientations, but not by a significant margin. Printing 0 degrees in the XY plane significantly increased the compressive and flexure strengths of the material compared to other orientations.IntroductionAs additive manufacturing is becoming an integral part of engineering processes and more parts are being printed for final use rather than prototypes, it is important to understand which build orientations give the part longevity depending on the parts specific use.

This paper focuses only on Fused Deposition Modeling (FDM) as there are many more 3D printing techniques. FDM is an additive manufacturing technology commonly used for modeling, prototyping, and production applications [1]. Figure 2 shows the FDM process, laying down material in layers using plastic filament or metal wire that is unwound from a coil and supplies material to produce a part [2]. Many parts can be built in various directions and orientations to achieve the same result. Typically, 3D printed parts are built using an orientation that will use the least amount of time and material. The question is, would this affect the strength of the part being built? In order to test this, samples were designed in CAD software to be printed in 5 different orientations with respect to the X, Y, and Z-axis using ABS P430 produced by Stratasys. The samples were designed based on ASTM standards corresponding to the type of mechanical test (tension, compression, and flexure).

The samples were designed based on ASTM standards corresponding to the type of mechanical test (tension, compression, and flexure). Figure 1 displays how each tensile specimen was oriented for printing. Compression and flexure samples were also oriented in this fashion.Build speed of the uPrint is closely related to material use. In general terms, a lesser amount of support will allow for greater build speed. Another factor affecting speed is the axis orientation. The printer can build faster across the XY plane, than it can along the Z-axis. Therefore,939orienting a part so that it is 'shorter' within the modeling envelope will produce a quicker build. A model is stronger within a layer than it is across layers.

Therefore,939orienting a part so that it is 'shorter' within the modeling envelope will produce a quicker build. A model is stronger within a layer than it is across layers. Depending upon what features a part must demonstrate, it may need to be oriented in a way to have its greatest strength across a specific area [3].Loyola Marymount University’s Mechanical Engineering Department has performed research studying the effects of varying layer orientations of ABS plastic using Fused Deposition Modeling. This study focused on ABS P400, also produced by Stratasys. In the area of tensile strength, samples printed in the 0 degrees orientation were significantly stronger (ultimate/yield) than those printed in any other orientation with an average ultimate strength of 20.6 MPa. The weakest orientation observed was printed 45 degrees with an average ultimate strength of 7 MPa [4].

The weakest orientation observed was printed 45 degrees with an average ultimate strength of 7 MPa [4]. ABS P400 is similar to P430, so we expect to see results like those seen in Table 1.Table 1: Tensile Properties of ABS P400 Based on Printing Orientation [4].In another study of the anisotropic material properties of fused deposition modeling ABS, there was not much of a difference in the compressive strengths of the samples based on build orientation. Although the compressive strength of FDM samples was higher than the tensile940strength, the ultimate and yield strength was not affected much by build direction. Because of the anisotropic behavior of the parts made by the FDM process, the strength of a local area in the part depends on the direction of the beads of material (roads) relative to the loading of the part rather than the angle at which the parts were printed.

Raster orientation and air gaps were most important in determining strength in the samples [5].ZYXFigure 1: Specimens in various degree orientationsMethods and MaterialsPrinting3D Printer: uPrint SE Pluso Model Material: ABS P430 o Support Material: SR-30 soluble o Layer resolution: 0.254mmuPrint SE Plus 3D printer has fixed parameters. Machine doesn’t allow changing any parameters.941Figure 2: FDM Process [2]TestingApparatus:The apparatuses used for these experiments are as follows:Testing Machine: MTS 810 Material Test System Load Unit  Drive Mechanism: MTS Hydraulic Power Unit  Load Indicator: MTS FlexTest System Electronics  Micrometers: Whitworth Digital Caliper o Range: 0.01-150mm o Resolution: 0.01mmFigure 3: MTS Servohydraulic Test System942Test Specimens:For tension, test specimens were manufactured according to ASTM D638.

Dimensions used corresponded to Type I dogbone specimens, having an overall length of 165 mm, an overall width of 19 mm, a depth of 7 mm, a narrow width of 13 mm, a narrow length of 57 mm, a radius fillet of 76 mm, and a gage length of 50 mm [6]. According to ASTM standard D695, the specimen dimensions for unreinforced plastics in compression are a cylinder of size 12.7 mm in diameter by 25.4 mm in length [7]. As for flexure, ASTM standard D790 states the dimensions for thermoplastic materials are to be 127 mm in length, 12.7 mm in width, and 3.2 mm in depth [8].Conditioning:All specimens were conditioned at room temperature, approximately 24°C, at least 40 hours prior to testing. Each mechanical test was performed under similar temperature conditions to conditioning [6] [7] [8].Number of Test Specimens:Normally, the standard number of isotropic test specimens per sample required by all three ASTM standards is at least five.

However, due to time constraints, only three specimens per sample were tested [6] [7] [8].Speed of Testing:The speed of each tension test was held at a constant rate of 5 mm/min [6], while the speed of each compression and flexure tests was held at a constant rate of 1.3 mm/min [7] [8], each relative to the motion of the test fixtures.ProcedureTest specimens were modeled in MasterCAM according to their respective ASTM standards. They were then exported and printed on the u-Print SE Plus 3D printer in ABS P430. For each mechanical test, three specimens were printed for each orientation on various axes: 0 degrees in the XY plane, 45 degrees in the XY plane, 90 degrees in the XY plane, 45 degrees in the Z plane, and 90 degrees in the Z plane.943ResultsTensionFigure 4: ABS P430 FDM Tension SampleFigure 5: Tension Stress vs. Strain Curves of ABS P430 in Various Orientations Tensile Stress vs.

This distance was decided based on the nature of ABS plastic. A preliminary test was conducted on a cylindrical sample. The sample was compressed until it was almost entirely flattened by the test fixtures. The data collected showed that there was no definite point of failure, so a cutoff point needed to be established, as seen in Figure 7 [7]. It should be noted that strain is computed as moving head displacement of moving head divided by gage length. Since no strain gages were used, strain and modulus data is only for relative comparison between various orientations.945FlexureFlexure tests were conducted as per ASTM D790 and sample is exhibited in Figure 8. Figures 9 display load vs moving head displacement curves for various orientations. It should be noted that displacement of moving head is assumed as deflction while computing flexural modulus.

Figures 9 display load vs moving head displacement curves for various orientations. It should be noted that displacement of moving head is assumed as deflction while computing flexural modulus. Since no LVDT was used, flexural modulus data is only for relative comparison between various orientations.Figure 8: ABS P430 FDM Flexure SampleFlexural Load vs. Displacement Curves for ABS P4302500-deg XY200N,d a o L15010045-deg Z90-deg Z45-deg XY5090-deg XY001234567Displacement, mm0-deg XY45-deg XY90-deg XY45-deg Z90-deg ZFigure 9: Flexure Load vs. Displacement Curves of ABS P430 in Various Orientations946DiscussionTable 2 indicates the tensile, compressive, and flexural test data for all categories. Values in parenthesis indicate standard deviations. Figures 10 and 11 show the strength and moduli of all five orientations, respectively.

Values in parenthesis indicate standard deviations. Figures 10 and 11 show the strength and moduli of all five orientations, respectively. ABS P430 printed in 90 degrees in the Z plane shows the highest tensile strength of the five orientations, equaling 10.8 MPa. This orientation also has the largest percent elongation and the lowest modulus of elasticity, 6.92% and 586 MPa, respectively. However, the tensile strengths of the other four orientations fall very close to the highest tensile strength recorded. The percent difference between the largest and smallest tensile strengths was calculated to be 14.3%, a relatively small margin. According to the manufacturer, ABS P430 has a tensile strength of 37 MPa, a percent elongation of 3%, and a tensile modulus of 2320 MPa [9]. The drastic difference in values between the collected data and the manufacturer’s data could be a result of a difference in printing methods or printing accuracy.

The drastic difference in values between the collected data and the manufacturer’s data could be a result of a difference in printing methods or printing accuracy. Regardless, this data can still be used to compare strengths amongst the different build angles. For compression, ABS P430 printed 0-degrees in the XY plane resulted in the largest compressive strength, compressive yield strength, and modulus values, 59.3 MPa, 44.1 MPa, and 854 MPa, respectively, compared to the other orientations. There was a 67.4% difference between the highest compressive strength of 0- degrees in the XY plane (59.3 MPa) and the lowest of 45-degress in the XY plane (29.4 MPa). There is no data currently available from the manufacturer on the compressive properties of ABS P430, but from comparing the data collected from these samples to one another, 0-degrees in the XY plane appears to be the strongest overall.

In flexure, ABS P430 printed 0-degrees in the XY plane resulted in a strain value of 5 mm/mm, and a flexural strength of 122 MPa, the highest of any of the orientations. ABS P430 printed 45-degrees in the XY plane resulted in a strain value of 5 mm/mm, and also a flexure modulus of 3630 MPa. The percent difference between the highest and lowest flexural strengths (122 MPa and 59.5 MPa) equaled to 69.2%, while the percent difference between the highest and lowest flexural modulus (3630 MPa and 2200 MPa) was 49.1%.

Although, ABS P430 90-degrees in the XY plane proved to have highest percent elongation of 6.92% but lowest tensile modulus. 90-degrees in the XY plane orientation shall provide better toughness. ABS P430 90-degrees in the Z plane had the highest tensile modulus of 737 MPa but lowest percent elongation. 90-degrees in the Z plane orientation shall provide makes material stiff but at the expense of toughness. Build angles seem to play a larger factor in the compressive and flexural properties of this material. A compressive strength of 59.3 MPa and a flexural strength of 122MPa, both orientated 0-degrees in the XY plane, were calculated to be the highest values in their respective categories.The highest compressive modulus was 854 MPa for 0- degrees in the XY plane whereas highest flexural modulus was 3630 MPafor 45-degrees in the XY plane.

atomical Modeling and Rapid Prototyping Assisted Surgical ReconstructionT. Jiang(1), F. Lin(1), S. I. Kaltman(2) and W. Sun(1) (1) Department of Mechanical Engineering and Mechanics, Drexel University, Philadelphia, PA 19104 (2)Dept. of Oral and Maxilloficial Surgery, Nova Southeastern University, Fort Lauderdale, FL 33328AbstractA CAD-based approach to generate three-dimensional anatomical modeling and its application in analysis and rapid prototyping for surgical reconstruction is presented. The process of modeling technique, starting from processing computed tomography segmentation profiles, reverse engineering to reconstruct 3D anatomical model, and rapid prototyping to fabricate physical models are described. Methods of applying the least-square fitting criteria to simplify CT or MRI anatomical data, to generate NURBS based curves, surfaces, and solids, and to develop a CAD-based anatomical modeling for femur are introduced.

Feasibility of using finite element method for femur structural analysis and using rapid prototyping to reproduce tactical medical model for surgical reconstruction is also discussed.1.IntroductionComputed tomography (CT) and magnetic resonance imaging (MRI) technology has been successfully used to provide true three-dimensional anatomical views for internal structures of the human body [1-4]. The application of this technology has now moved forward with the utilization of CAD/CAM and rapid prototyping technology (RP) to allow a rapid fabrication of the physical tissue models or custom-made prostheses from CT or MRI data [5-8]. RP produces extremely detailed models that serve as excellent templates for the creation of custom implants, offers the surgeon a three-dimensional tactile imaging modality, and display local regions of interest, such as surgeon to draw round a tumor on the CT image and have it built into the model for disease diagnosis in the prototyping anatomic model.

Both computer anatomical model and RP physical mode have been utilized as an accurate planning tool for reconstructive surgeons of many specialties, covering from oral surgery, plastic and reconstructive surgery, neurological surgery, orthopaedic surgery, and craniomaxillofacial surgery [9-12].Commonly used anatomical modeling reconstructions are usually based on the voxel representation. The anatomical images are developed through either contour generation, or surface extraction, or volumetric representation. Contour-based method generates 3D-like displays from sliced CT or MRI outlined segmentation profiles of tissue interested [13,14]. 3D surface extraction uses the full three-dimensional nature of the tomography data to directly produce a geometric surface description [15]. 3D volumetric representation directly produces the appearance of 3D surfaces without to explicitly define their geometries.

3D volumetric representation directly produces the appearance of 3D surfaces without to explicitly define their geometries. Although those models are capable of describing the anatomical morphology and are used to rapid prototyping through a converted STL format, due to the voxel-based nature, those models may not suitable for customized design and anatomical structural analysis.

For example, the contour generation represents only surfaces without the internal details, 3D surface extraction requires either a large amount of computational power or extreme sophistication in data organization and processing, and 3D volumetric model, while produce a realistic 3D anatomical appearance, does not contains geometry and topological relations in the database of the model [16, 17].The objective of this paper is to present our study on developing a computer-aided modeling technique for anatomic tissue reconstruction, and its application in modeling analysis and rapid prototyping for reconstructive surgical planning. Activities involving in our modeling approach, starting555from processing computed tomography imaging and CT or MRI 2D segmentation profiles, using reverse engineering to reconstruct 3D anatomical model, and using rapid prototyping to fabricate physical model will be presented.

Details of the modeling algorithm to convert CT or MRI anatomical data into a CAD- based solid model, and to use the model for analysis and simulation will be described.

Application in rapid prototyping to reconstruct tactical medical model for surgical reconstruction is also discussed.2.Framework of developing a CAD-based three-dimensional anatomical modelingThe CAD-based three-dimensional anatomical modeling is developed based on using the commercial available medical imaging processing software (MIMICS), a computer-aided design software (Pro/Engineer), and the B-rep CAD technique for generating NURBS based curves, surfaces and solids.CT ImagesSegment CT ImagesReverse Engineering SoftwareConvert to ContourSTL fileRapid PrototypingProfile PolylineNURBS-based Contour/Surface Solid Modeling ConverterSegment ProfileNormalized Profile PointsConstruct Flank SurfaceAdd Top and Bottom SurfaceGenerate Solid Model3D Solid ModelMultiple ApplicationFEM Model and Generate elementsApply Boundary ConditionRun FEM AnalysisBiomechanical Analysis ResultsFigure 1: Framework for developing CAD-based anatomical modelingThe framework for developing the CAD-based anatomical modeling is outlined in Figure 1.

By creating a computer-generated contour following a single CT or MR image intensity threshold value of the interested anatomy, we start to generate NURBS curves based on the cloud point data exported from556MIMICS’ sliced CT or MRI segmentation profiles. A series of ordered NURBS curves are generated from the 2D segmentation profiles and those curves are collected to generate NURBS surfaces and solid in a CAD (Pro/Engineer) environment. Thus generated surface or solid contains the built-in geometry and topological relations in the modeling database. Particular attention is paid to generate interface for anatomic structures with bifurcated protrudes.3.Approach to construct CAD based anatomic modeling3.13D segmentation to contour polygonsThe interface from CT or MRI scanner to rapid prototyping has already been well developed. For example, Figure 2 shows a bony image of femur represented by the 3D STL data generated from MIMICS.

For example, Figure 2 shows a bony image of femur represented by the 3D STL data generated from MIMICS. However, this “solid” image doesn’t contain the real geometry and topologic relationship for the represented solid. In general, most medical image processing software are not capable to introduce the CAD-based data structure into the model database because of their voxel-based nature. Some software may provide interface between the medical modeling and CAD modeling, however, they are difficult to be used in modeling complicated anatomical structure, particularly for anatomic structures with bifurcated protrudes, such as a femur model shown in Figure 2.Figure 2: Femur anatomic structure (3D STL representation)To develop a CAD-based model for femur structure shown in Figure 2, we start with obtaining the polygons exported from MIMICS at each segmentation layers (Figure 3a).

Those polygons are constructed as geometrical contours formed by the points on the segmentation profiles (Figure 3b). Those polygons are then fitted through the least square criteria to reduce the number of the points interpolated in the NURBS curves (Figure 3c). Closed surfaces from the NURBS curves are then generated and a B-rep solid model can be created from these surfaces.(a) femur by contour representation (b) polygons from MIMICS(c) fitted polygons557Figure 3: Contour polygons3.2Polygon approximation by the least square fitting methodThe curves shown in Figure 3b are zigzagged because there are too many cloud points on the profiles. It is not necessary to take all the vertices as interpolation points to generate NURBS curves. So the least square fitting method is used to reduce the number vertices on the polygons before the NURBS curve generation. The closed polygon (profile) is divided into segments before the fitting.

So the least square fitting method is used to reduce the number vertices on the polygons before the NURBS curve generation. The closed polygon (profile) is divided into segments before the fitting. There are two alternatives in this curve segmentation:1)Dividing the polygon according to the number of points, for example, given a particular number of points as one fitted segment. This approach is easy to be implemented, but it is not adaptive to handle complicated geometry.2)Dividing the polygon according to the distance. The algorithm for this approach is as following: First, we calculate the total distance of the closed contour, dist, which is the summation of the distance of each segment. Second, we select a segmenting number num according to precision required. The larger of the number num is, the more accurate of the fitting. Each fitted segment is determined by: dist/num.

Second, we select a segmenting number num according to precision required. The larger of the number num is, the more accurate of the fitting. Each fitted segment is determined by: dist/num. This approach is adaptive to the shape of the closed contour so it is used in this model reconstruction.The least square criteria can be represented as:f(= (cid:140) )[2 iD(i2 ])(1)where, axis, f( ) is a function of can be calculated by dist/num. Schematic view of these parameters is shown in Figure 4.is the angle between the fitting line and X axis,i is the angle between each segment and X , Di is the length of each segment, and D is length of the goal segment whichFigure 4: Polyline fitted by least square criteriaThe anglecan be determined from solving following derivation equation:ƒf(/)=ƒ0(2)The fitting algorithm is written in LISP program.

The fitted polygons are shown in Figure 3c.5583.3Generating NURBS curvesIn order to generate a NURBS surface, the contour NURBS curves need to be generated first. The sequence of control points will have an influence on the shape of the generated NURBS surface. As shown in Figure 5, two sets of closed curves have the same control points, but the sequences of control points are different, so the generated surfaces from those curves are different. The first surface is generated by sorted curves (Figure 5a), the second one is generated from the original sequence of points with unsorted end points and the direction. So it creates a twisty surface (Figure 5b).

Hence, it can be solved by Gaussian elimination without pivoting. Thusgenerated NURBS curves for the femur structure are shown in Figure 6.559Figure 6: Femur model represented by NURBS curves3.4Generating NURBS surface and solid modelingpoints {The global interpolation of the surface can be constructed based on a set of (n+1) x (m+1) data lkQ , },,...,0=,...,0=knmmandQ, lk=vuS k(,l)=nm( NuN), pik, qj() Pv , ji l(6)i=0j=0in which,lv are parameters, andjiP,are the unknown control points. Since the contour curves arealready generated,jiP,can be computed more efficiently through the sequence of curve interpolations.For a fixed l Equation (6) can be expressed as:Q, lk=n( uN , pik)(mN, qj() Pv , ji l)=n( RuN , li), pik(7)i=0j=0i=0R , li=mN, qj() Pv , ji l(8)j=0=,, kQ lk,...,0liR , are the v =.

If we fix i and let l vary, Equation (8) lv 0, ,..., 0, ,..., RnEquation (7) represents the curve interpolation through the points.control points of the isoparametric curve on S(u, v) at fixedR iP iP , miwill represent the curve interpolation through the points, withas the computed, micontrol points. Therefore, the control points can be obtained by using knot vector and parameters, and the . Figure 7 shows three 4th- curve interpolation can be performed through=,, kQ lk,...,00, ,...,nR iRand, midegree NURBS surfaces with a C2 continuity surfaces generated by above curve interpolation.560Figure 7: Generated three NURBS surfacesIt should be noted that the femur structure has bifurcated protrudes. Most modeling software are unable to generate the single surface for bifurcated structure because there are two closed contours in each scanning layer beyond the bifurcated section.

Most modeling software are unable to generate the single surface for bifurcated structure because there are two closed contours in each scanning layer beyond the bifurcated section. Therefore, an interface layer is generated in order to connect the bottom diaphysis and top two bifurcated protrudes. This interface layer is shown in Figure 8.An important property of NURBS curve is the local approximation, as explained by Les Piegl iW is changed, the curve is only affected at . Based on this property, we divide the control points on the the portion on the interval original curve into two intermediary curves. By adding new control points, we generate two 4th-degree NURBS curves, curve 1 and curve 2, as shown in Figure 8. These two new curves are geometric identical to the bifurcated section curve. With these two intermediary curves, two new bifurcated surfaces can be generated and added to the top two protrudes to form the closed surfaces.

With these two intermediary curves, two new bifurcated surfaces can be generated and added to the top two protrudes to form the closed surfaces. With the interface layer as the surface for diaphysis, three solid models can be generated and then merged to form a single solid model as shown in Figure 9.iP is moved, or the weight [18]. That is, if the control Point [ , i uuºu]1++ piCurveCurveBifurcated sectionFigure 8: Intermediary NURBS curves for bifurcated sectionFigure 9: A single solid model for femurFigure 10: Mesh and BC of bone FEA model4.Application in finite element analysis and rapid prototyping4.1Integrated CAD and finite element analysisThe CAD-based femur model is applied for finite element analysis. Three-dimensional mesh of the bony model is generated from Pro/MECHANICA. The modulus of elasticity of the bone is taken from the assumed relationship between bone elastic modulus and apparent density presented by Carter and561Heyes [19].

The modulus of elasticity of the bone is taken from the assumed relationship between bone elastic modulus and apparent density presented by Carter and561Heyes [19]. The mesh representation, loading and boundary condition is shown in Figure 10. Predicted stress and strain distribution is shown in Figure 11 and Figure 12, respectively.Figure 11: Von Mises stress distributionFigure 12: Strain distribution in the boneIt should be pointed out that the objective of current finite element analysis is to explore the (Pro/Engineer) with CAE-based analysis feasibility of (Pro/MECHANICA), and to test the model capability of automatic meshing generation, rather than to study the detail of biomechanical properties of the femur structure. Therefore, only preliminary FEA results are presented.integratingthe CAD-based model4.2Application in rapid prototypingApplications of using rapid prototyping to produce physical anatomic models are presented in Figure 13 and Figure 14.

Figure 13 shows a prototyped hip anatomy prototyped by 3D Printing (Z-corp). The prototyped major bones included the upper portions of the left and right femur, the pelvic bones, and the lower end of the spine (sacrum and coccyx). The model includes the entire pelvic region with all of the bones connected and keeps the proper anatomic position between them.Fig. 13: Prototyping model of hip anatomy Fig. 14: Prototyping model of skull with tumorFigure 14 shows a rapid prototyping model of a skull with a tumor. The growth of the tumor is visible and has created a large protrusion from the skull. In order to analysis the volume and anatomic position of tumor, the tumor identification is necessary.

The growth of the tumor is visible and has created a large protrusion from the skull. In order to analysis the volume and anatomic position of tumor, the tumor identification is necessary. Both anatomic model and prototyping model of tumor and skull without tumor are constructed and the prototypes are shown in Figure 15.562Figure 15: Rapid prototyping model with tumor identification5.ConclusionA CAD-based modeling approach to construct three-dimensional anatomical modeling is presented. Procedures of the modeling approach and applications of the model in structural analysis and rapid prototyping for surgical reconstruction are given. The presented computer modeling approach is useful to model bifurcated anatomic structure.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperTHE ANISOTROPIC YIELD SURFACE OF CELLULAR MATERIALSKaitlynn M Conway*; Zachary Romanick*; Lea M Cook*; Luis A Morales*; Jonathan D Despeaux*; Marcus L Ridlehuber*; Christian Fingar*; Daquan Doctor*; Garrett J Pataky* *Department of Mechanical Engineering, Clemson University, Clemson, SC 29631AbstractMechanical metamaterials are often limited in engineering applications because of uncertainty in their deformation behavior. This uncertainty necessitates large factors of safety and behavior assumptions to be included in mechanical metamaterial designs, detracting from the largest benefit of metamaterials: their ultralight weight.

In this study, a yield envelope was created for both a bending dominated and a stretching dominated cellular material topology to improve the understanding of the response of cellular materials under various load types and orientations. Experimental studies revealed that the shear strength of a cellular material is significantly less than that predicted by the Mohr’s criterion, necessitating a modification of the Mohr’s yield criterion for cellular materials. Both topologies experienced tension-compression anisotropy and anisotropy dependent on the topology orientation during loading with the stretching dominated topology experiencing the largest anisotropies.A substantially improved version of this paper appears in a special issue of the TMS publication JOM, March 2022.IntroductionMonolithic materials, in terms of specific strength, are limited by their elemental weights creating a ceiling of what can be achieved via solid materials alone.

An increasingly popular method for manipulating a material’s properties, is through the geometric arrangement or topology of the material’s mesostructure [1]. Low density metamaterials combine solid materials and zero density voids to occupy material property spaces not achievable by monolithic materials [2]. The complex geometric freedoms additive manufacturing (AM) affords enables for the creation of metamaterial topologies with significantly altered mechanical properties compared to the base materials [3]. However, understanding of the mechanics responsible for these metamaterial’s deformation behavior, particularly fracture and failure, across different environments and loading situations is understudied, necessitating large factors of safety [1].

This conservative approach detracts from the largest benefit of metamaterials: their ultralight weight.Low-density metamaterials have evolved from honeycomb sandwich boards [4] and disordered foams [5] to highly complex topologies across a variety of size scales and base materials [1]. The rapid advances in metamaterial use and design has resulted in the introduction of new topologies to the scientific community, outpacing the understanding of failure mechanisms. Additionally, there is not a consistency among studies of base material, cell size, or AM process which makes comparing23various topologies across reported studies difficult at best and speculative at worst. Mechanical metamaterials are known to have tension and compression asymmetry [6] as well as orientation-dependent anisotropic behavior when loading due to lack of rotational symmetry of many topologies [7].

Additional studies of metamaterials have found the shear strength of metamaterials to be significantly less than that of the uniaxial strength [8]. Due to this, metamaterials will have a complex, non-trivial yield envelope that is crucial for engineers to understand to design metamaterial components with confidence.The goal of this study was to determine the effect of topology, particularly stretching and bending dominated topologies, on the yield surface size and shape. Mohr’s criterion was chosen for this study as it was designed to handle tension-compression anisotropy and the failure envelope can be determined with three tests: tension, compression, and pure shear [9].Methods and MaterialsFour cellular metamaterial topologies of interest were AM by fused deposition modeling: a traditional honeycomb, an auxetic honeycomb, a triangular, and a circular topology.

Topologies were modeled with a 5 mm unit cell and a 0.6 mm wall thickness and were printed on a Dremel 3D45 printer using Dremel Eco-ABS white filament [10]. Each topology was tested in two orientations, deemed σ1 and σ2, where σ2 is a 90° in plane rotation of σ1, Figure 1.Specimens of all four topologies were tested in tension, compression, and shear to populate the Mohr’s Circle. Three specimens of each type were tested in each loading condition and the yield strengths calculated by a 0.2% offset were averaged. Novel pure shear grips previously designed by the authors were used to insure pure shear rather than a combined shear and uniaxial loading state [8].Figure 1: Cellular material topologies tested with original orientation (σ1) and 90° rotation (σ2)24Experiments were performed in displacement control at a rate of 30 μm/s using a hydraulic load frame for the shear specimens and a screw-driven load frame for the tension and compression samples.

A single camera captured a 6-by-4 unit cell cross section of the center of each specimen during tests and digital image correlation was used to calculate the strains in the cellular material during deformation. The camera used was a Point Grey model GS3 equipped with a Navitar lens and a 1X adaptor and captured images at a rate of 1 Hz. Images were used to calculate the full field displacement and strains using commercial digital image correlation software, VIC 2D. Strain calculations were performed with a virtual strain gage of 21 pixels2 and a spatial resolution of 35 pixels2 following the procedure outlined in [11]. Global material properties were calculated using the area methodology outlined in [7].ResultsShear StrengthExperimentally determined tension, compression, and shear yield stresses for each orientation of the chosen topologies were plotted as a Mohr’s circle, shown for the honeycomb topology in Figure 2a,b.

The max shear line is calculated by plotting a linear line between the maximum shear value of each Mohr’s circle. The max shear line is displayed in green in Figure 2a,b. The max shear line was selected rather than a tangential line to the Mohr’s circles offering a more conservative yield envelope. As cellular materials are sensitive to defects from the AM process, a conservative yield criterion was preferable. Using the max shear line failure envelope, a yield surface was created for the orientations, shown in Figure 2c [9]. The two yield surfaces were collapsed to one surface for the material using all six experimentally determined yield strengths to create a surface showing the full tension-compression-shear anisotropy as well as the orientation specific anisotropy, Figure 2d.

The yield surfaces for each orientation are shown in Figure 3, where the top row is the yield surface for each orientation, the middle row is the combined yield surface, and the bottom row is an image of the corresponding topology.and𝜎𝜎1𝜎𝜎225Figure 2: Honeycomb Mohr's Circle and Yield SurfaceFigure 3: Yield Surface for selected topologies. Top row: Yield surface for each orientation. Middle row: Full field surface for topology. Bottom row: Selected topology. Scale adjusted for the triangle orientation26The decreased shear strength of these cellular materials resulted in a discontinuity at the pure shear loading condition. The metamaterial was weaker in shear than would be expected from the maximum-distortion-energy theory or the more conservative maximum-shear-stress theory [12]. This behavior was seen in every topology.The stretching and bending topologies have a topology-driven strength-ductility trade off, Figure 4a.

This behavior was seen in every topology.The stretching and bending topologies have a topology-driven strength-ductility trade off, Figure 4a. Stretching dominated topologies were stronger, while bending topologies were more ductile. This has been well-documented in metamaterial research [13–16]. In the stretching dominated triangle topology, there is a high concentration of strain very localized at the strut junctions, even at low global displacements. These strains are circled in red in Figure 4b, taken at 0.5 . With further loading, the triangle topology yields and then catastrophically fails when a crack develops at one of the highlighted junctions. In the bending dominated topologies, the honeycomb and auxetic unit cells in Figure 4c and Figure 4d respectively, as the metamaterial is loaded, stress is primarily concentrated in the vertically aligned struts. This concentration of stress causes localized buckling of a strut, circled in red in Figure 4c.

This concentration of stress causes localized buckling of a strut, circled in red in Figure 4c. When a strut buckles, the stresses on the neighboring struts increase and the neighboring struts buckle as well creating a row-by-row collapse of the metamaterial. Figure 4c,d were taken immediately after yield to illustrate the buckling cell wall.𝜎𝜎𝑦𝑦𝑦𝑦𝑦𝑦𝑦𝑦𝑦𝑦Figure 4: a) Shear stress-strain curves of selected topologies b) Strain maps of triangle in shear c) Strain maps of honeycomb in shear d) Strain map of auxetic in shear27Topology EffectsThe stretching dominated triangle topology has the largest yield surface with a maximum tensile, compressive, and shear strength of 21.1 MPa, -28 MPa, and 12 MPa, respectively, over the honeycomb topology, which had the second largest yield surface, with a maximum tension, compressive, and shear strength of 12 MPa, -11.9 MPa, and 7.4 MPa, respectively.

The triangle topology has a different scale than the other topologies in Figure 3 due to the greater strength. The increased strength of the honeycomb topology in comparison to other bending dominated topologies is due to the uniform deformation of the honeycomb unit cells during loading and the uniform formation of plastic hinges at the sharp corners, circled in red in Figure 5a. As the honeycomb is deformed, plastic hinges form as the arms rotate, absorbing energy and increasing the toughness. Uniform deformation allows the plastic hinges to develop without the stress concentration at any one plastic hinge becoming so large the metamaterial prematurely fractures. Figure 5 is shown after yielding to better illuminate the locations of strain and plastic hinges within the topologies, but the onset of crazing at plastic hinges has been shown to occur prior to yielding [7,18]. The circle topology does not have sharp corners for plastic hinges to form.

The circle topology does not have sharp corners for plastic hinges to form. Additionally, the hourglass shape of each strut creates a local stress concentration in the middle of the length of the strut identified by the high strain circled in red in Figure 5b. With increased loading, the local stress concentration increases leading to early fracture of individual struts lowering the yield stress, creating a smaller yield surface than the honeycomb topology. The auxetic honeycomb also develops plastic hinges when loaded; however, the auxetic honeycomb has a negative Poison’s ratio [19]. As the auxetic honeycomb is loaded in tension, the horizontal struts are loaded in compression due to the negative Poison’s ratio (the struts in the honeycomb topology are loaded in tension). The auxetic unit cells rotate when loaded to decrease the compressive stress on the horizontal strut, changing the loading path and non-uniformly increasing the stress concentrations at the plastic hinges.

The auxetic specimen yields when the plastic hinges with the largest strain concentrations, such as that circled in Figure 5c, fracture.Figure 5: Strain maps of topologies in theorientation under tensile loading, all figures the same scaleTopological-Induced Anisotropy𝜎𝜎1The triangle topology had the largest orientation anisotropy with an 80% change in tensile strength and a 66% change in compressive strength from loading in the orientation. The circle topology had the next largest orientation anisotropy with 25% tensile and 𝜎𝜎2 27% compressive difference in strength. When the triangle topology is loaded in the direction the axial load is primarily distributed among a third of the struts, the struts parallel to theorientation to the𝜎𝜎1𝜎𝜎128direction of loading. The positive axial load path is marked in red in Figure 6a and supported by the concentration of positive strains along these struts shown in Figure 6c.

The positive axial load path is marked in red in Figure 6a and supported by the concentration of positive strains along these struts shown in Figure 6c. When loaded in the direction, the axial load is distributed among the angled struts, Figure 6b, consisting of two thirds 𝜎𝜎2 of the struts in the topology as demonstrated by the positive strains in Figure 6d.

This difference in load distribution between the orientations explains why the tension and compression yield strength of the orientation is nearly twice the tension and compression yield strength of the orientation.𝜎𝜎2Figure 6: Axial load in a stretching dominated topology is distributed along the struts parallel to the loading direction, in the triangle topology in the orientation the axial load is distributed primarilyamongst a third of the struts, in the𝜎𝜎2orientation the axial load is distributed amongst two thirds of the𝜎𝜎1 strutsDiscussionThere is a debate about whether to treat a mechanical metamaterial as a continuous material and use a homogenization analysis [20] or a structure and analyze the behavior using the local behavior driven by the unit cell topology [21]. Both approaches have merit and are particularly useful in specific applications.

Both approaches have merit and are particularly useful in specific applications. During elastic deformation, the homogenization approach of considering the metamaterial as a uniform material simplifies identifying global material properties, regardless of local plasticity that is occurring before fracture. However, identification of local stress and strain accumulation within the unit cell throughout loading is essential for understanding and predicting differences in the plastic and fracture behavior between topologies. Some studies have incorporated a two scale approach to incorporate both methodologies when analyzing29𝜎𝜎1metamaterials to use the local behavior of the lattice to determine the global behavior of the material while limiting the computational work required to analyze an entire lattice structure [14,22].

In this study, a global evaluation of the metamaterial was used to calculate the yield strength of the selected topologies; however, analysis of local strain concentrations within the unit cells were investigated to explain the change in yield surface shapes between topologies. Local analysis of the lattice was not limited to a single unit cell but rather the 6-by-4 unit cell cross section captured by the camera where digital image correlation was performed. This consideration of several unit cells rather than a single cell allowed for evaluation of uniformity and the measurement of unbalanced local deformation such as the rotation and uneven plastic hinges in the auxetic topology.

Evaluation of stress concentrations and local deformations within the lattice unit cells were then connected to the metamaterial yield surface size and shape to analyze how topology drives the yield surface.The poor shear strength of cellular materials creates a yield surface shape that would not be expected from a monolithic material due to the formation of local plastic hinges and strains non- uniformly distributed among the metamaterial’s struts. Yielding of the metamaterial is caused by buckling of a cell wall or fracture of a cell wall or joint on the mesoscopic scale rather than at the microstructural level as would occur with a monolithic material. This departure from traditional mechanisms of yielding in a material is shown by discontinuities on the yield surface that occur at the pure shear loading. The large variation in the size of the yield surface of the tested metamaterials further supports the need to consider metamaterials, at least partially, as a structure.

The large variation in the size of the yield surface of the tested metamaterials further supports the need to consider metamaterials, at least partially, as a structure. All topologies in this study were produced using the same filament material, yet the maximum compressive strength of the triangle topology was over three times that of the auxetic topology and twice the honeycomb topology. Additionally, the extreme strength anisotropy between was amplified beyond a tensile-compressive strength differential, but also orientation dependence due to changing loading paths. This behavior is more typical of a structure than a conventional material.ConclusionA yield surface was created for four cellular material topologies tested in tension, compression, and shear. The yield surfaces demonstrated cellular materials were weakest in shear loading and have large anisotropies dependent on loading condition and orientation.

The yield surfaces demonstrated cellular materials were weakest in shear loading and have large anisotropies dependent on loading condition and orientation. Local strain fields were used to identify the mechanisms driving the deformation of different topologies and the driving effect of the unit cell deformation mechanism on global plastic properties. Buckling of cell walls and the formation of plastic hinges were the main mechanisms to cause yielding within the materials; however, rotation of the cell walls due to its negative Poison’s ratio caused the auxetic topology to have the smallest yield surface.AcknowledgementsThe authors would like to thank Student Pathways in Engineering and Computing for Transfer Success, NSF Grant 1834081, at Clemson University for support of this research.30References[1] X. Yu, J. Zhou, H. Liang, Z. Jiang, L. Wu, Mechanical metamaterials associated withstiffness, rigidity and compressibility: A brief review, Prog. Mater. Sci.

isotropy in Alumina Processed by SLS.P. Kamatchi Subramanian, N.K. Vail, lW. Barlow and H.L. Marcus Center for Materials Science and Engineering, The University of Texas at Austin.Abstract Alumina powders of 15~m size and 2~m size were processed by SLS using PMMA and a copolymer. The 2~m powders were agglomerated and mixed with the polymer powder before being processed by SLS. SLS bend strength specimens were made with parts built along different orientations. The variation of the strength with incident energy density and with orientation was studiedIntroductionSelective Laser Sintering (SLS) of ceramics is very challenging because of the high melting points of the ceramics. In SLS one of the operating parameters, the energy density is defined as the amount of laser energy input per unit area 1 in each layer. In the case of alumina to form lines of sintered material energy densities of the order of 1000 callcm2 is required.

In the case of alumina to form lines of sintered material energy densities of the order of 1000 callcm2 is required. Even at this high energy densities the parts formed do not hold together because of the huge thermal gradients. One solution is to use a low melting temperature second phase material which will melt under the laser and bond the particles together. This second phase may be inorganic2,3,4 or organic5,6. This paper describes some of the SLS experiments of alumina with organic binder. This paper focuses on the variation of properties of the green shapes with processing parameters including the orientation of scanning in the SLS process.ExperimentsTwo kinds of alumina powder were investigated in this study. They were i) 2~m powder from Golden Technologies and ii) 15~m powder from Norton Corp. SLS of 2~m as received powder was unsuccessful using a polymer binder even at compositions of 35 Vol % polymer.

They were i) 2~m powder from Golden Technologies and ii) 15~m powder from Norton Corp. SLS of 2~m as received powder was unsuccessful using a polymer binder even at compositions of 35 Vol % polymer. In an attempt to overcome this problem it was decided to agglomerate the particles. Through a series of experiments of heating to various temperatures for varied lengths of time it was decided that 1400e for 3hrs would give a desirable agglomeration. The 2~m powders were heated in a kiln to 1400e in 12 hrs at a constant heating rate and held at 1400e for 3hrs. The powder was also cooled down in about 12hrs. This produced agglomerates of an average size of 1OO~m. The flow diagram is shown in Fig.l.The organic binders used were i) PMMA and ii) a copolymer, Poly[methylmethacrylate co-butyl methacrylate] (80 mol%MMA/20 mol%BMA). The preparation of these polymers is explained elsewhere7.

The preparation of these polymers is explained elsewhere7. The PMMA had a melt index of 31.5g/10 min and the copolymer had a melt index of 11.6g/10min as 111easured following ASTM D 1238 at 200e with an extrusion pressure of 0.52 MPa (75psi).3302JJ.m alumina powder 88 receivedPolymer emul810n PMMA or CopolymerAgglomerationI) Heat gradually to 14000 II) Hold at 14000 for 3 hrs III) Gradual CooIdownSpray drying emulsion to obtain PMMA powder or Copolymer powderMix agglomerates and polymer powder~SLSFigure 1. Processing of 2Jl1D Alumina powderi1BuildDirectioni iFigure 2. Orientations used in building the bend strength specimens.

Processing of 2Jl1D Alumina powderi1BuildDirectioni iFigure 2. Orientations used in building the bend strength specimens. i:build along thickness, scanning along width; ii: build along thickness scan along length; iii: build along width with scanning along length; iv: build along width, scan along thickness 331The 15Jlm alumina powders were spray dried with the copolymer in a Anhydro spray drier with inlet temperature of 200C and outlet temperature of 80C. The agglomerates of 2Jlm alumina were mixed with the polymer powder obtained by spray drying the polymer emulsion. Two mixtures were made, i) the agglomerates of 2Jlm alumina mixed with the spray-dried PMMA powder and ii)the agglomerates of 2Jlm alumina mixed with the spray-dried copolymer powder. The spray-dried and mixed powders were then SLS processed.The layer thickness was kept constant at 175Jlm.

The spray-dried and mixed powders were then SLS processed.The layer thickness was kept constant at 175Jlm. The laser scanning speed and laser scan line spacing were varied along with the power of the laser to change the incident laser energy density. The scan line spacing was varied between 75Jlm and 125 Jlm and the scan speed was varied from 30cm/s to 150 cm/s. The laser power was varied between 6 and 14W. The energy density was varied between 2 cals/cm2 and 8 cals/cm2. Three point bend strength specimens 0.076mxO.025mxO.00625m (3 I x1 I xO.25") were made by SLS. The specimens were made in 4 orientations i) built along thickness with scanning parallel to width ii) built along thickness with scanning parallel to length, iii) built along width with scanning parallel to thickness and iv)built along width with scanning parallel to length as shown in Fig.2.Results and DiscussionSLS of parts from both the spray-dried and mixed powders produced strong green parts.

For the case of the spray-dried 15Jlm powder built along thickness with scanning parallel to width, condition (i), the variation of density with incident energy density is as shown in Fig.3. The density varies little with incident energy density. The variation of bend strength of the green parts with incident energy density is shown in Fig.4. It may be seen that the green strength increases initially with energy density because of better melting and wetting of the alumina by the molten polymer. At still higher energy densities the strength decreases due to decomposition of the polymer.Scanning along the length while building along thickness, condition (ii) produces a lower green strength at lower energy densities than those scanned along width and built along thickness, condition (i). This anisotropy is due to the region scanned by the laser having longer time for heat transfer between two consecutive scans for condition (ii).

This anisotropy is due to the region scanned by the laser having longer time for heat transfer between two consecutive scans for condition (ii). At higher energy densities Le., around 6cals/cm2 the strength of the samples made by scanning along length has a higher strength due to degradation of polymer for condition (i). Even in the case of scanning along length we have a decrease in strength at higher energy densities. But this decrease occurs at still higher energy densities.With increase in energy density, higher temperatures are reached in the sample being made. As temperature increases well above the glass transition temperature, the degradation of the polymer increases. Temperatures in the decomposition regime are attained at lower energy densities in the case of samples with shorter scan lines because the time between successive scan In the case of samples made with longer scan lines the temperatures in the overlays is smaller.

Bend strength of samples from lSJ.lI11 alumina coated with copolymer built along thickness.333~.- a. ""'" ... .c C) C ... CD a- enU)... -- .c... c.l= ...... ~fIj Q.('J.)2502001500 00 00 0condition ivIcondition iiiI§0100~t:I501234 5 Energy density (cal/cm2)6Figure 5. Strength of samples from 15J.l.m alwnina coated with copolymer built along width.775" 0\5500 §condition ii000 0D08 83250D §Ocondition ii_~__-.i~,_ .._.;_~_,....L_J1007.7J1015.5 3.25 Energy density (c a II c ) Figure 6. Strength of samples from agglomerated 2J.Un alumina mixed with PMMA built along thickness.334densities the strength decreases for reasons stated before. Once again it may be seen that when the scan lines are longer the green strength obtained at lower energy densities is lower. Again the peak in the strength vs.

Once again it may be seen that when the scan lines are longer the green strength obtained at lower energy densities is lower. Again the peak in the strength vs. energy density plot occurs at a higher energy density for the parts made with longer scanlines.SLS of agglomerated 2J1m alumina mixed with 30 V% PMMA.Specimens for bend-strength were made from this mixture in the four orientations described above. For specimens built along thickness, conditions (i) and (ii), Fig.6 it may be seen that the maximum strength is higher than that obtained for the 151lm alumina spray-dried with copolymer. This is due to the fact that the alumina agglomerates are much larger than 151lm and hence have a lower surface area to be wet by the polymer. The polymer content is also slightly higher. The variation of green strength with energy density follows similar trend as in the previous case with the strength increasing with energy density at lower energy densities. The peak in the strength vs.

The peak in the strength vs. energy density plots occurs at a higher value of the energy density for the parts made with longer scan lines.When specimens were built along the width at low energy densities the parts made with shorter scan lines, condition (iv) had higher green strength, Fig.? A slight increase in density with increase in energy density may also be observed, Fig.8.SLS of the mixture of agglomerates of2J1m alumina with 20 V% copolymer.Specimens for bend strength measurement were made from this mixture by building only along the thickness direction. Three different scanning methods were employed. They were i) scanning along width, ii) scanning along length and iii) scanning alternate layers in length and width. The density increases slightly with increase in energy density, Fig.9. Scanning along the length produces samples of lower density compared to the other two modes at lower energy densities. In this case also it may be seen (Fig.

Scanning along the length produces samples of lower density compared to the other two modes at lower energy densities. In this case also it may be seen (Fig. 10) that the strengths are lower for scanning along the length at low energy densities, as described earlier. The case of mixed (alternate) scanning results in samples of strength intermediate between the two other modes. This is to be expected since alternate layers will have higher and lower strengths corresponding to scanning in the other two modes.SummarySuccessful SLS was done on 21lm alumina particles by agglomerating the powders and mixing with polymer powder prior to Laser Sintering. Samples were built in different orientations in the SLS process. In all cases the samples that were built using shorter scan lines had a higher strength at lower energy densities.

Samples were built in different orientations in the SLS process. In all cases the samples that were built using shorter scan lines had a higher strength at lower energy densities. For both longer and shorter scan lines the green strength increases initially with increase in energy density, attains a peak and then starts decreasing with further increase in energy density due to polymer degradation. The peak in the case of samples built with longer scan lines occurs at a higher value of the energy density due to lower temperatures reached under those conditions.3355condition iv4,-.... eu ~ ~ ......., ..c ...... CI) 2c <» a- ..... rJ:J30 0 00 § 00 00 8 000~ 000/ condition ...0lii1o1 23456Ener2Y Density (cal/cm2)7Figure 7.

KLE FOOT ORTHOTICS OPTIMIZATION BY MEANS OF COMPOSITE REINFORCEMENT OF FREE-FORM STRUCTURESJ. Munguia*, KW. Dalgarno *School of Mechanical and systems engineering. Newcastle University, Newcastle, UKAbstractAnkle-Foot-Orthoses (AFOs) are assistive devices prescribed for a number of physical and neurological disorders affecting the mobility of the lower limbs. Traditional AFO design is based on thermoforming polymer sheets around a plaster cast with a characteristic stiffness depending on the selected trim-line with the inherent design limitations. Additive Manufacturing has been explored as an alternative process; however it has proved to be inefficient cost-wise.

Additive Manufacturing has been explored as an alternative process; however it has proved to be inefficient cost-wise. This work explores the possibilities of generating modular AFO elements: calf, shank, and footplate with localized composite reinforcement that aids in the optimization of the device in terms of functionality, aesthetics, rigidity and cost.1.IntroductionAnkle-foot orthoses (AFOs) are orthopedic devices usually prescribed to address a number of common walking impairments such as stroke, multiple sclerosis or cerebral palsy [1, 2]. Their main purpose is to compensate the malfunctioning of the neuromuscular system which manifests in an unhealthy gait pattern localized in the foot and ankle e.g. drop foot, weak ankle and other impairments.

drop foot, weak ankle and other impairments. The term “orthoses” refers to medical devices applied externally in order to help patients to gradually acquire healthier gait patterns.Different types of AFOs are normally prescribed for various impairments depending on the intensity, range of movements and patient complexion. Some typical designs include: leaf spring AFOs, hinged, with dorsal cover, etc. [1]. Traditionally, patient-specific AFOs are made by one of two methods: 1) thermoforming copolymer sheets of a prescribed thickness (normally 3-4mm) around a plaster-cast impression or 2) Prepeg carbon fiber sheet layup using a plaster cast base as a reference surface. The typical plastic AFO manufacturing process is shown on figure 1 [3].a) Figure 1. Typical stages for PP AFO manufacturing.

The typical plastic AFO manufacturing process is shown on figure 1 [3].a) Figure 1. Typical stages for PP AFO manufacturing. a) Cast rectification b) PP sheet placement c) Edge cutting d) Add-onsb) c) d )Carbon-fiber based AFOs offer stiffer and lighter structures and are manufactured by laying up thin layers of carbon fiber on a rectified positive casts, with a predefined bonding agent (epoxy resin), and alternating their orientation until the desired thickness / stiffness is achieved.Typical productivity for both types of AFO manufacturing processes (plastic and carbon fiber) depends on manual labour skills (Table 1). Observed average production rates for a single worker on a per-day basis are 1- 3 carbon fiber AFOs and 2-4 polymer vacuum formed AFOs [3]. Lead times for both types of AFOs can vary significantly; while for polymer- based AFOs it is typically 1 week, for carbon-fiber AFOs it will vary between 1 to 3 weeks until the final delivery.

Lead times for both types of AFOs can vary significantly; while for polymer- based AFOs it is typically 1 week, for carbon-fiber AFOs it will vary between 1 to 3 weeks until the final delivery. The main parameters dictating the AFOs performance are the material type, thickness, trim line (lateral carving) [3-5], localized reinforcement, and footplate properties [1] which are normally selected by personal preference of the orthotist.Table 1. Cost and productivity target for AFOs [3]Product cost Typical productivity Total lead time€170 - €220 each 3-4 full AFOs per working day Approx. 10 days to reach customer2. Additive Manufacturing of AFOs Given the artisanal manufacturing style of personalised AFOs, additive manufacturing has been identified as a natural method to produce automated highly personalized devices that fit every patient’s specific requirements [3].

Work on additive manufacture of AFOs has focused on highlighting the freedom of design associated with additive technologies. Pallari et.al [4] demonstrated the design, manufacturing and optimization possibilities of an SLS- made AFO scanned from a preexisting Polypropylene device. Schrank et.al [6] analyzed the dimensional accuracy of passive dynamic ankle-foot orthotics made by SLS by manufacturing scaled-down versions of the device, while Mavroidis et. al. [7] presented an AFO customization process by manufacturing two samples with the sthereolithography process which showed similar properties to a conventional PP AFO. Muraru et.al. [8] presented a different approach by modeling the necessary mechanical properties for a finite element material definition of laser sintered AFO devices. Other related research is based in new design-freedom and customization capabilities of additive manufacturing from a pure- design perspective [6, 8, 9].

Other related research is based in new design-freedom and customization capabilities of additive manufacturing from a pure- design perspective [6, 8, 9]. However none of the previous research has addressed the feasibility of AM as a dedicated method for AFO manufacturing particularly concerning two factors: economics and performance optimization. 2.1 The build-time and cost limitation of Additive manufacturing for AFOs One problem not usually described in literature is the build time and cost required for producing a full size AFO in one piece by any of the AM technologies. Even for the most versatile of processes such as SLS, obtaining a high productivity compared to conventional technologies can be impractical. A number of build time estimations for AFOS have been calculated for different product volumes.

A number of build time estimations for AFOS have been calculated for different product volumes. A generic 3D AFO design (Figure 2) has been placed in two SLS machines and a second stage of optimization has been performed by minimizing the Z height and packing multiple components in a single build. Results are presented on Table 2.Figure 2. Generic AFO 3D STL file and SLS-made AFO Duraform PATable 2 Build time estimation for generic AFO designAlthough build-time optimization can be pursued by making an efficient use of the build space the software-based build-time estimations suggested Ankle Foot Orthotics are not well suited for full size manufacture due to the high build time and costs involved, including material waste, machine preparation and generally lower productivity when compared to well established manufacturing methods.

One approach for optimizing additive manufacture of AFO devices is the design of modular/independent elements which can me joined together by further secondary operations such as localized reinforcement with composite materials. The use of composite materials such as carbon fiber/grass fiber- epoxy has been recurrent for automotive, aerospace and lightweight structures design and a number of cases make use of additive manufacturing structures as sacrificial structures used for the molding of composite materials [10, 11]. This research explores the generation of modular/individual structures than can be joined into a single full-sized component by the use of localized composite reinforcement.

This research explores the generation of modular/individual structures than can be joined into a single full-sized component by the use of localized composite reinforcement. For this study a low cost system (Bits from Bites- FDM) printing in PLA material was used for both: assessing the mechanical properties of localized reinforcement and obtaining a prototype orthotic device.the use of additive manufacturing for3 Method 3.1 Flexural tests specimen preparation A total of 30 test specimens were prepared in compliance with [12] for a 3-point bending test in order to quantify the flexural properties of the PLA material along the 3 main build directions (X, Y & Z). Specimens were produced in a Bits from Bytes (low cost-FDM machine) and the build prepared using the Axon v2.0 software with the following parameters: layer thickness of 0.25mm, speed multiplier 1.3X, fill density 100%, fill pattern: linear.

Specimens were packed in batches of 5 and the three groups were produced with a raft-base layer for further stability, which was fully removed prior to the tests. Fifteen specimens (5 for each orientation) were tested in plan PLA material while the remaining 15 were subjected to a localized reinforcement by applying two layers of carbon- fiber/polyester fabric (East Coast Fiberglass, UK) with a combination of epoxy-hardener equivalent to a 5:2 mixing ratio respectively. A preliminary layer of epoxy+ hardener agent was applied to the upper surface of the specimens and then after each successive carbon-fiber -layer. Specimens were left to cure for a period of 36 hours and then the excess hardened fiber was shaped to follow the original rectangular shape of the specimen.3.2 Evaluation of mechanical properties of plan FDM specimens The prepared specimens were tested in a Tinius Olsen H25KS universal testing machine with a 5k N load cell.

The test setup showed on figure 3 followed the test procedure described by BS-EN-ISO 178:2003 with a cross sectional speed of 5mm/min and an automatic displacement limit of 50mm. The first set of experiments was performed on plain PLA specimens and then second batch on the reinforced specimens with the composite material facing downwards.Figure 3. 3-point bending of PLA + reinforcement specimen3.3 Design of modular complex shapes for an AFO A prototype AFO device requested by a professional orthotists (Peacocks Medical Group, Newcastle, UK) was developed which is intended to measure the counter-force ejected by patients with hemiplegia. Some hemiplegic patients tend to develop a gait-deficit exhibited in the form of a knee hyperextension pattern while walking. Thus the proposed design would allow a full dorsiflexion movement while inducing a resistance to plantar flexion that is then recorded by a load cell placed at a given distance from the heel joint.

Thus the proposed design would allow a full dorsiflexion movement while inducing a resistance to plantar flexion that is then recorded by a load cell placed at a given distance from the heel joint. The prototype AFO device was designed in a modular way so that all elements fit the available build space on a BfB 3D touch machine. The full design was developed in Autodesk Inventor v13. and the file conversion/preparation was performed on the Axon v2.0 machine software. The effect of orientation was neglected as the full assembly was not intended to withstand loads during walking, but to act as a receptacle for localized reinforcements with a carbon fiber composite. The AFO elements requiring reinforcement were prepared on the relevant surface by removing all trace of support structures and leaving a plain smooth surface. A formulation of 5:2 Epoxy resin/hardener was applied to the base surface and two layers of carbon fiber sheet applied.

A formulation of 5:2 Epoxy resin/hardener was applied to the base surface and two layers of carbon fiber sheet applied. The set of components was left to cure during a period of 36 hours at room temperature to allow for a complete cure cycle.4 Results 4.1 Flexural tests results The comparison of the flexural properties of plain PLA specimens against carbon-fiber reinforced under the 3-point test is shown on figures 4 and 5. The localized reinforcement with 2 layers of carbon-fiber and a 5:2 epoxy/hardener ratio provided an increase in flexural strength of 50% average, split along different orientation axis as follows: X=40%, Y=47%, and Z=71%.140) a P M(120100t h g n e r t Sla r u x e l F806040CF-reinforcedPlain -PLA200xyzFigure 4.

Flexural strength for both: plain- PLA and PLA + CF reinforcement along 3 build orientationsFor the flexural modulus an increase of nearly 30% average was registered which was also consistent along the 3 main orientations: x=33%, y=22.6%, Z= 29% (figure 5).40003500) a P M3000(s u u d o Ml25002000CF-reinforcedPlain PLAla r u x e l F150010005000xyzFigure 5 . Flexural strength for both: plain- PLA and PLA + CF reinforcement along 3 build orientationsAdditionally the presence of a carbon-fiber reinforcement layer on the lower surface prevented the specimens from breaking during the test as they did when reinforcement is not applied.

The major bending stresses concentrate on the lower face at the middle span as opposed to a conventional 3-point bending tests where compressive stress on the upper surface equals the tensile stress on the lower surface (figure 6).Figure 6- PLA specimens after test with a higher stress concentration area on the lower surface interfacing the carbon-fiber.4.2 Design of modular complex shapes for an AFO A prototype AFO device for the measurement of knee hyper extension forces was designed in Autodesk Inventor v13. Figure 7 shows the concept where various AFO parts were modeled separately and joined as an assembly so they can be placed individually in the build chamber of low cost FDM equipment (figure 8).Calf elementMain strutLoad cell socketAnkle jointFoot plateFigure 7. 3D CAD of AFO jig for in-lab stiffness measurement with footplate, calf and post section printed in PLA (Blue) and Carbon fiber reinforcedFigure 8.

3D CAD of AFO jig for in-lab stiffness measurement with footplate, calf and post section printed in PLA (Blue) and Carbon fiber reinforcedFigure 8. FDM fabrication of multiple components for FDM- carbon fiber reinforced AFOThe total build time for the separate elements was 16 hours, however various trials were tested where the full set of components could be embedded in one single build cycle below 12 hours, mainly by increasing the layer thickness parameters and deposition rate. The resulting parts in PLA material were individually reinforced by means of carbon fibre/Blue polyester (East Coast Fibre glass, UK) plus epoxy resin in order to provide extra rigidity in selected areas: foot plate/external surface, main strut/both sides. Results are shown in figure 9 where the final structures comprise a prototype stifness measuring jig.Figure 9.

Results are shown in figure 9 where the final structures comprise a prototype stifness measuring jig.Figure 9. Final FDM-made AFO jig fabricated by multiple PLA components - carbon fiber reinforced (UNEW)5 Discussion5.1 Conventional vs Additive manufacturing of AFOs Althought ankle foot orthotics have been studied by previous research they remain a challenge for AM-based production due to the overall build-size and the time implications of building each device in one piece. The productivity of conventional methods still surpases that of AM for both, SLS and FDM however the advantages of digital fabrication lay on the design oportunities when designing the final devices, as no plaster casts are required to mould the plastic. A modular design approach has been introduced where the three main elements of an AFO: Calf, footplate and strut are split and arranged within the build envelope so the full set of components is produced in one single build and assembled afterwards.

Due to the need for support structures the FDM method is not the best technology for optimizing packing ratios as the rotation of individual components to minimize the z-height has a direct affect on the build time due to the need for support structures (figures 10 and 11).Therefore the optimum orienttaion for FDM is when parts can be built without the need of support structures.Figure 10. Build time for calf part built along the Z axis in Axon v2.0 software.Figure 11. Build time for calf rotated 90 degress to minimize Z height via support structures.5.2 Effect of localized reinforcement with composite material The localized reinforcement of sacrificial structures has been studied in different fields from automotive applications [11] to complex tooling design [10] or art artifacts [13]. However the specific effect that a predetermined composite material formulation can provide to a structure made by additive manufacturing had not yet been reported.

However the specific effect that a predetermined composite material formulation can provide to a structure made by additive manufacturing had not yet been reported. As the results presented in this work suggest, a localized reinforcement of two layers of carbon fibre-polyester with a 5:2 epoxy-hardener ratio have an effect of nearly 50% increase in flexural strenght and approximately 30% in flexural modulus across the 3 main build directions.

However the most important characteristic is the structural estability the composite provides to strenghten the AM-made components preventing rupture even after the plastic section has failed.Further work could address the effect of different epoxy/hardener ratios, the use of additional reinforcement layers or the localized reinforcement on both sides of the was out of the scope of this work this work suggests the use of low cost additive manufacturing parts as composite-deposition surfaces or guides can fin applications in end-use products that do not necessarily need an oversized plastic structure but a lighweight-complex overal shape with good mechanical properties.6. Conclusion The use of low cost AM components in conjunction with composite materials can open new avenues for the design of high complexity products without the need for expensive tooling components or secondary operations such as plaster cast, wood mould machining orthermoforming substrates.

However, plastic components being used as a guide-surface or sacrificial skeleton must provide sufficient rigidity for the composite deposition stages, particularly when pressures will be applied by means of vacuum baggin, resin transfer moulding, or simply by applying uniform presures manually.The solution presented in this work for an orthotic device is based on thin walled components printed along the Z axis without the need of support structures. It has been observed that adhesion of PLA to CF-polyester composites is sufficient for prototyping purposes however additional tests must be performed acording to the final purpose of the device, e.g. epoxy resin penetration range vs. wall thickness, tensile tests of hollow vs dense polymer substrates, etc. This work introduces the low-cost 3D printing plus reinforcement approach as an alternative route for the design and manufacture of orthotic devices with complex shapes and localized reinforcement needs.

It is expected that new applications add-up to increase the body of knowledge about the behaviour of such products which will mix both areas, composites theory and additive manufacturing.AknowledgementsThis work was funded through the European Commission Framework Seven Program (grant number NMP2-SE-2009-228893) as part of the A-Footprint project (www.afootrint.eu).References 1- Daan J.J. Bregman 2011 The Optimal Ankle Foot Orthosis The influence of mechanical properties of Ankle Foot Orthoses on the walking ability of patients with central neurological disorders. ISBN: 978-90-6464-486-3 2- Tina Taiming Chu 2001. Biomechanics of Ankle-Foot Orthoses: Past, Present, and Future. Top Stroke Rehabil 2001;7(4):19–28 3- A-FOOTPRINT 2013- Ankle and foot orthotic personalization via Rapid Manufacturing. FP7 Programme, Grant Agreement Number: NMP2-SE-2009-228893. http://www.afootprint.eu/ 4-Jari et al.2010. Design and Additive Fabrication of Foot and Ankle-Foot Orthoses.

The AM process requires a layer-by- layer execution to build parts, allowing for a unique opportunity to collect data and monitor the process in real/semi-time. The incipient phase of AM monitoring and control typically consists of developing an automated unsupervised statistical anomaly detection algorithm that is capable of detecting irregularities through parameter measurement and sensing features. In this paper, we develop a simple and effective method for detecting anomalies through use of statistical distances from data collected during the laser-based Directed Energy Deposition (DED) AM process.IntroductionConventional manufacturing techniques have long been used to produce parts for a variety of applications however, the present-day push to produce parts with complex shapes in a time- competitive and sensitive market has led to great interest in metal AM.

The AM process has completely revolutionized the field of manufacturing because of its ability to Computer streamline Aided Designs (CAD) to produce parts that reachFigure 1. Schematic of FormAlloy L2 metal AM system with MPS Camera1606the market in reduced time, reduced waste when preparing parts, ability to utilize multiple materials during production and print harsh complexity demands. Figure 1 highlights the DED process used to produce parts for this investigation. DED is a metal AM method that involves the use of thermal energy to melt material during deposition (Tofail, 2018). During the DED process, metal powder (or wire) is fed from the feedstock until it reaches contact with the laser of the electron beam and is melted. Then it is deposited on top of a substrate, this process repeats until the final layer of the part is printed (Zhang, 2018).

Then it is deposited on top of a substrate, this process repeats until the final layer of the part is printed (Zhang, 2018). The DED technique is widely used among researchers and industry professionals because it produces high-quality parts with a high degree of controllable grain structure (Tofail, 2018). In this investigation, we use a Laser-based Directed Energy Deposition (LB-DED) technique (Shown in Figure 1) with a powder flow method for the production of metal parts. There has been a tremendous amount of research completed to identifying and classifying defects in metal AM parts (Roh, 2021) however, many of these techniques are computationally time-consuming or require ex- situ methods to identify the defects.

This investigation serves to present the idea of utilizing an unsupervised statistical monitoring technique with real printing features as an initial phase of Quality Control (QC).MethodologyTwo hollow cylinder samples were produced, made of Inconel 625 and were fabricated by FormAlloy using a FormAlloy L2 metal AM system, shown in Figure 1 (FormAlloy 3D Metal Printing Machines and Technolo, 2022). The L2 system works in a similar fashion to how most metal AM processes function. First, the powder nozzles permit the powder stream to continuously flow, supplying powder particles to the laser beam. Next, the stream and particles come into contact with the laser beam which melts the particles and deposits the new material onto the substrate. The entire method ensues under the supervision of (MPS) the Melt Pool Size camera. The L2 system uses a laser as the main proponent of the heat source in contact with metal the feedstock.

The entire method ensues under the supervision of (MPS) the Melt Pool Size camera. The L2 system uses a laser as the main proponent of the heat source in contact with metal the feedstock. This system comes fully equipped with a continuous wave (CW) fiber laser utilizing a focused spot diameter of 1.2 mm and maximum output power of 2 kW. sensing component is an MPS camera that was set in the path of the laser. As the melt pool moves, the camera measures the width of the pool during the entire The printingData-driven AlgorithmOutput/Post processInput/In-situStatistical Anomaly DetectionCTSensingpowderfromProcess ParametersThein-situprocess.1607uniqueness of the L2 system is that it provides users with real-time measurement of the melt pool dynamics.

In total, two samples were produced, one low-quality (sample 1) and one high-quality (sample 2).Figure 2 highlights the layout of this investigation, which began with a visual analysis of the printed parameters collected in real-time. Next, a feature selection is used to truncate the original dataset and an unsupervised statistical method is utilized for outlier detection of the selected printing features. Finally, ex-situ processing takes place to identify and measure defects within the parts. Acceptable and unacceptable layers were identified through a user adjustable technique. Literature in the research community defines feature selection as the act of reducing a subset dataset from an original complete feature dataset through the use of a feature selection technique (Cai, 2018). Modern forms of data are measured and recorded at extremely high rates however, not all data is deemed informative.

Modern forms of data are measured and recorded at extremely high rates however, not all data is deemed informative. When data is highly dimensional or dynamical (or both), it benefits the experimenter to reduce the useable dataset, especially when the correlation among features is highly. When feature selection is executed, many times the overall computational time during the experiment is reduced and model performance can sometimes be increased. The Random Forest algorithm is widely accepted by literature for a feature selection technique due to its interpretability and ability to provide the user with important values (Petkovic, 2018). The RF technique works to reduce variance through the bagging (bootstrap aggregation) technique. It comprises a mean of noisy and unbiased models (Hastie, 2009). The main aspect of RF is the tree-growing process where arbitrarily selected input variables are bootstraped on the data.

It comprises a mean of noisy and unbiased models (Hastie, 2009). The main aspect of RF is the tree-growing process where arbitrarily selected input variables are bootstraped on the data. The purpose of this is to de-correlate the trees without swelling the variance (An introduction to feature selection, 2013). RF chooses . The main process is driven by the classification predictor, mathematically represented as (Hastie, 2009):Figure 2. Design Diagramnumber of variables by randomly pulling fromfeatures so that𝑚𝑚𝑝𝑝𝑚𝑚 = √𝑝𝑝𝐵𝐵 (𝑥𝑥) = 𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚 𝑣𝑣𝑚𝑚𝑚𝑚𝑣𝑣��𝐶𝐶̂𝐵𝐵(𝑥𝑥)��1 (1) . Figure 3 Where highlights the result of the two samples produced during printing and the corresponding printer features collected in-situ during the printing process.𝐵𝐵 𝐶𝐶̂𝑟𝑟𝑟𝑟is the total number of trees that we are predicting on a new point𝐵𝐵𝑥𝑥1608Figure 3.

Sample 1 and 2 (left) and Inter-layer low-density identification (right)To determine if the condition of a part has deviated significantly from its normal state, we highlight the use of a statistical method for identifying outliers. For this multivariate number of observations. Figure is a visual dataset, we define explanation of how the technique identifies outliers. The Mahalanobis squared distance 𝑛𝑛 has been utilized to detect multivariate outliers for structural health applications and used in this investigation. It a non-negative scalar defined as (Farhidzadeh, 2014):number of features and𝑝𝑝2𝑇𝑇−1= (𝑥𝑥 − 𝜇𝜇)⋅ (𝑥𝑥 − 𝜇𝜇)𝐷𝐷⋅ 𝐶𝐶Where, correspond to the mean vector and covariance matrix, respectively.is a vector measurement that corresponds to a possible outlier,and𝑥𝑥𝜇𝜇𝐶𝐶Results and DiscussionBefore advanced detection techniques are used, simple unsupervised methods may present a possible solution to quick and reliable outlier detection.

Based on a feature selection using RF (importance values shown in Table 1), five features are focused on for further statistical analysis. This step is crucial to limiting the useful information collected during the AM build process. Figure 4 presents the mean, maximum, minimum, and standard deviation trends of the printing parameters. It serves the purpose of an initial visual stage of identifying any peculiarities during the metal AM build. Several outliers can be seen in the Laser Power feature, Figure 4b. Some minor outliers can be visualized from the MPS feature in Figure 4a. The other features do not demonstrate any clear trends for a user to visually identify.Table 1. Feature selection results FeatureImportance Value1609(2)0.191 Laser Power 0.181 Melt Pool Size Shield Gas 0.113 Powder Carrier Gas 0.121 0.033 Feed RateFigure 5 highlights the Mahalanobis Distance results of both samples. In Figure 5a, the low-quality sample shows three distinct levels of measurements.

In Figure 5a, the low-quality sample shows three distinct levels of measurements. This is not the case in Figure 5b, the high-quality sample showing only two levels of distance values. The low- quality sample shows a high frequency of squared Mahalanobis Distance values between 3.5-7, where the high-quality sample showcases the highest frequency of Mahalanobis Distance values between 4.5-4.8. It is clear from the calculations of the squared Mahalanobis Distance that a distinguished separation of multivariate measurements takes place among high/low-quality samples.a)b)c)d)1610e)Figure 4. Visual statistical analysis of a) MPS, b) laser power, c) feed rate, d) sheild gas, and e) powder carrier gasX-Ray Computed Tomography (CT) was performed on both samples to provide porosity information. The distribution of porosity diameter and the number of porosity per layer are shown in Figures 6 and 7.

The distribution of porosity diameter and the number of porosity per layer are shown in Figures 6 and 7. A deeper inter-layer analysis for samples 1 and 2 in Figures 6 and 7, respectively. Squared Mahalanobis Distance values are able to identify a high that frequency of porosities cause defects. For sample 1 (Figure 6b), of ‘unacceptable’ layers based on an adjustable is highlighted in the blue shaded region and corresponds to the Mahalanobis Distance values calculated in Figure 6a. The high-quality sample (sample B), shown in Figure 7, highlights the unsupervised statistical anomaly method to detect maximum porosity values (Figure 7a). Visual clarity is present between maximum porosities early in the build (outlined with purple) and later in the build (outlined in red).is shownahighfrequencyex-situthresholda)2𝐷𝐷Layer Numberb)2𝐷𝐷Layer NumberFigure 5.

The unique layer-by-layer process of AM provides an exceptional opportunity to monitor and control the quality of parts. This study demonstrates the usefulness and effectiveness that a simple unsupervised anomaly detection method can have on early defect detection and prevention. Included in the study is a feature selection technique and visual analysis that aid in this incipient phase of the manufacturing process. If on- boarded to an AM system, this technique could give a user confidence while monitoring the build process and potentially prevent defects without the need for advanced defect classification methods.References1. Cai, J. L. (2018). Feature selection in machine learning: A new perspective. NeurocomputingNeurocomputing, 70-79.2. Farhidzadeh, A. D.-N. (2014). Post-earthquake evaluation of pipelines rehabilitated with cured in place lining technology using acoustic emission. Construction and Building Materials, 54, 326-338.3.

plication of a Diagnostic Tool in Laser Aided Manufacturing ProcessesSashikanth Prakash, Mallikharjuna Rao Boddu and Frank Liou Department of Mechanical, Aerospace and Engineering Mechanics University of Missouri - Rolla Reviewed, accepted August 19, 2003 AbstractLasers play a vital role in producing parts with high dimensional accuracy, strength and quality in today's rapid prototyping industry. In the process of Laser Metal Deposition, many problems are encountered where the part quality does not meet the required standards. This could be due to the nonconformity of control parameters or unnecessary interactions between the control factors. This paper discusses the implementation of Dr. Genichi Taguchi’s optimization techniques using Design of Experiments (DOE) where a series of fractional factorial experiments are performed on the laser deposition process.

Genichi Taguchi’s optimization techniques using Design of Experiments (DOE) where a series of fractional factorial experiments are performed on the laser deposition process. The results from these experiments are evaluated with respect to the rate of deposition alongside the part quality and the optimized level setting of control parameters are determined efficiently. This tool can be used to detect and diagnose flaws and discrepancies in the Laser Metal Deposition process and optimize it accordingly.IntroductionLaser Metal Deposition (LMD) is a process where metal powder is focused on a substrate and a laser beam melts the powder and deposits it. It is possible to convert any part into a series of slices and each slice can be deposited by the above method accurately (layered deposition) and the whole part can be fabricated.

It is possible to convert any part into a series of slices and each slice can be deposited by the above method accurately (layered deposition) and the whole part can be fabricated. The advantage of this process is that complex geometries can be constructed with high degrees of accuracy to achieve near net shape with a solid model of the part. The Laser Aided Material Processing system (LAMP, shown in figure 1) at the University of Missouri Rolla is primarily comprised of a 2.5 KW Nd: YAG Rofin Sinar laser (at TEM00) with integrated 5-Axis FADAL CNC with a maximum spindle speed of 7500 RPM.Figure 1. A Schematic of the Laser Aided Material Processing system (LAMP).There is a coaxial screw feed dual powder delivery system that is capable of delivering two types of powder for Functionally Gradient Material parts. The 5-axis CNC ensures high part finish quality and can be used to remove any unwanted deposit.

The 5-axis CNC ensures high part finish quality and can be used to remove any unwanted deposit. A schematic of the factor interrelationship is shown in figure 2.254This process has the capability of producing overhang parts and does not use support material because additional use of support increases the build time of the part and requires a time consuming post-processing. With a five-axis deposition integrated with five-axis machining, these obstacles are removed [1]. This paper deals with the testing of H13 powder deposited on H13 substrates and optimized for maximum build rate and at the same time, for having a good part quality. The hardness and the porosity will be used to evaluate part quality in this case study. There is a large range of layer thickness as well as deposition rates that can be achieved using laser deposition. The deposit rate can be increased by increasing the laser power, powder flowrate and the traverse speed.

The deposit rate can be increased by increasing the laser power, powder flowrate and the traverse speed. However, the requirement of a good part quality puts a limit on optimal deposition speeds. Both the layer thickness and the volume deposition rates are affected predominately by the specific energy and powder mass flow rate. Here, specific energy (SE) is defined as:SE = P/(Dv), (1)Where “P” is the laser beam power, “D” is the laser beam diameter and “v” is the process traverse speed. Also it is well known that actual laser power absorbed in the melt pool is not the same as nominal laser power measured from a laser power monitor due to absorptivity and other plasma related factors depending on the materials [2]. The use of adjusted specific energy is thus preferable.

The use of adjusted specific energy is thus preferable. Considering the factors, it has been reported that there is a positive linear relationship between the layer thickness and adjusted specific energy for each powder mass flow rate [3].Bead WidthBead HeightPart QualitySpecific EnergyPowder FeedrateCooling RateSpot SizeLaser PowerTraverse SpeedPorosityLaser Lens Focal LengthOverlapLaser Beam ModeTensile, Yield Strengths and HardnessDilutionDelaminationStand-off DistanceMicrostructureGrain Size and AlignmentEffect of Carrier GasFigure 2. Schematic showing the basic laser deposition factor interrelations.255Design of Experiments - Taguchi Methods of OptimizationThe Taguchi system of quality engineering is aimed at improving the robustness of a product and reducing the sensitivity to external variation. In other words, the part or product should perform as per customer expectations in the presence of various operational conditions.

In other words, the part or product should perform as per customer expectations in the presence of various operational conditions. A special set of arrays called Orthogonal Arrays (also known as fractional factorial experiments) is used to optimize the processes. Not all the combinations of parameter levels are used to find out the best setting for a particular result, but a well defined combination that spans the whole set evenly is used. The main advantage is that an optimized result can be achieved in the least possible time.Experimental ApproachThe steps that were used in the implementation of the Taguchi methods are:(cid:120)Identifying the factors and interactions: A total of 7 factors were chosen in this process as shown in table 1.(cid:120)Identifying the levels of each factor: It was found that 3 factors were necessary to check the linearity in the variation.

Moreover, there is a good chance of getting closer to the best values more quickly.(cid:120) Selecting an appropriate orthogonal array (OA): An orthogonal array “A” is defined as an “N * k” array with entries from the set “S” with a number of levels “s”, strength “t” and index “(cid:540)”[4]. The arrays must be chosen according to the number of control parameters, their respective levels and the factors which interact with each other. The L18 was chosen for this experiment.(cid:120) Assigning the factors and interactions to columns of the OA: One column was allocated to measure the factor interactions and any other factors and errors that were not taken into consideration.(cid:120) Conducting the experiments: Three samples were used for each experiment for repeatability. Analyzing the data and determining the optimal levels – this is an important part of the whole process and helps in obtaining the optimized values depending on the output we choose to measure.

Analyzing the data and determining the optimal levels – this is an important part of the whole process and helps in obtaining the optimized values depending on the output we choose to measure. In the analysis of data, Signal to Noise ratios (S/N) are used in controlling the target as well as reducing the variation about the target. The three types of Signal to Noise Ratio approaches are Nominal the Best, Smaller the Better and Larger the Better. Analysis of Variance (ANOVA) is used in this process is to calculate the percentage contribution of the control factors associated with the conclusions.(cid:120) Conducting the confirmation experiment: This is done to validate the accuracy of the experiment. If the predicted values that we obtain after optimization match the results from the confirmation experiment, then the experiment is valid.256Factor Level SelectionTable 1.

If the predicted values that we obtain after optimization match the results from the confirmation experiment, then the experiment is valid.256Factor Level SelectionTable 1. Control factor levels and their values chosen for this experiment.In order to determine the levels of the factors in the experiment, a series of experiments was conducted and it was ensured that there was a deposition and good bonding with the substrate. It can be noted that the inner gas pressure has only 2 levels and the other factors have 3 levels (table 1). The Degrees of Freedom (DOF) that we need to conduct for this experiment is 13. So it is appropriate to use the L18 orthogonal array for the experiment as shown in figure 4. Three samples were made for each experiment for repeatability hence the overall DOF for the control factors changes to 39. The main goal is to improve the deposit build rate and also maximize hardness (maximum of 120 in the Rockwell B scale) and minimize the porosity.

The main goal is to improve the deposit build rate and also maximize hardness (maximum of 120 in the Rockwell B scale) and minimize the porosity. These characteristics are combined into a single yardstick known as an Overall Evaluation Criterion (OEC) which is to be used to weight the output of the various combinations of the factors. The different values are first normalized and a weight is allotted to each characteristic to ensure a single standard is followed throughout the experiment. It is important to note that the deposits were four layers and had a square area with zigzag deposition pattern for subsequent layers.Table 2. The actual experimental layout of the L18 array.257The above setup (table 2) is a typical L18 array. It should be noted that there is a separate column attributed to the errors. This error column was added mainly to take into account the interactions and unaccountable noise factor effects that would affect the performance of the system.

This error column was added mainly to take into account the interactions and unaccountable noise factor effects that would affect the performance of the system. The main emphasis was to reduce the effect of the errors or variations and make the system more robust to these error variations and produce the expected result even in the presence of these errors. The array was set up and quality characteristics for the various combinations of control factors and their various levels were calculated and then the mean, standard deviation, and signal to noise ratios for the noise as well as the control factors were computed. The signal to noise ratio for the “Larger the Better” case was considered for the build rate and the hardness.

The signal to noise ratio for the “Larger the Better” case was considered for the build rate and the hardness. The three quality characteristics were then combined to form a single Overall Evaluation Criterion (OEC) and used for comparison and further calculations.Formulation of the Overall Evaluation CriterionIt is quite common to encounter a situation where there is more than one Quality Characteristic (QC). It is possible to analyze the QCs separately, but it cannot be guaranteed that a single common optimized result will be achieved. It would more often be different and contradicting. This emphasizes the use of what is called an Overall Evaluation Criterion. In this case a certain relative weight percentage must be allocated to each QC subjectively during the experiment planning session. The different QCs are adjusted in such a way that they give rise to a QC that is either Smaller the Better (STB) or Larger the Better (LTB).

The different QCs are adjusted in such a way that they give rise to a QC that is either Smaller the Better (STB) or Larger the Better (LTB). To combine the different QCs into a single OEC, they must be normalized first and then weighted. The Nominal the Best (NTB) QC must be modified to represent the deviation from the nominal value which would become STB [5]. In this case the build rate, hardness and the porosity were given a weight percentage of 40, 40 and 20 respectively. To simplify the experimental array, the individual sample readings have not been shown in the experiment table (table 3). There were three readings taken for each quality characteristic on each sample. Only their mean values are shown in the table. The OEC is formulated as a LTB case for all three characteristics. The hardness was measured in the Rockwell B scale, the build height and width were measured in an optical microscope and the time was calculated from the NC code for the deposition.

The hardness was measured in the Rockwell B scale, the build height and width were measured in an optical microscope and the time was calculated from the NC code for the deposition. The build rate was calculated as build volume per unit time in mm3/s. The porosity was viewed at 500X resolution and measured by a ranking system from a scale of 1 to 10. Here the OEC was formulated as follows [5]: (cid:173) (cid:176) (cid:174) (cid:176) (cid:175)QC QC (cid:16)QC QC (cid:16) a(cid:189) (cid:176) (cid:190) (cid:176) (cid:191)(cid:189) (cid:173) (cid:176) (cid:176) (cid:14) (cid:190) (cid:174) (cid:176) (cid:176) (cid:191) (cid:175) The above equation is for a larger the better case where QCa is the performance value of “a”, QCb is the performance value of “b”, QCmax & QCmin are the best and the worst performance values of the respective QCs and Wx is the weight percentage allotted to the QC “x”.

For NTB approaches, it must be noted that the OEC will be evaluated based on the magnitude of deviation from the nominal value. This magnitude must be made to be smaller the better. This can be easily reformulated to fit larger the better as shown above [5].258ExperimentTrialIGPFPF OG DO INTAvg.Avg.Avg. OEC 1 - 40:40:20 OEC 2 - 33:34:33No.

The experimental array with the measured performance statistics and S/N ratio.From table 3, we can see that there is an attempt made to evaluate another OEC with 33:34:33 weightage. This was done to study the change in the optimal values when all QCs are given equal importance. The OEC 2 can be used only if the part quality is of primary importance. It is also important to note that the optimal level settings may not be the same for different QCs. The optimal combination for hardness may differ from that of porosity. When both are used, there would be a trade-off in the result obtained.Analysis of the Means (ANOM)The Analysis of the Means (ANOM) is conducted basically to do a two step optimization. This is done by first reducing the variation in the process (using the S/N ratio and maximizing the slope) and then shifting the mean or target performance to get the best output. The ANOM was conducted next and the factor level plots were constructed.

The ANOM was conducted next and the factor level plots were constructed. The levels of the control factors with the highest S/N ratio were used in verification of the predicted values for the confirmation experiment [6]. The ANOM tables for the mean and the Signal to Noise ratio are shown below (tables 4 and 5).Mean 1 2 3Inner Gas Laser pressure Power WattPSI57.80 54.9442.84 60.1266.15Feedrate Powder Outer gasSpotOverlapFlow rate pressure Diameter FactorIPM52.6458.30 58.17g/min40.81 56.9771.33PSI58.31 54.44 56.36mm58.77 57.10 53.24% 47.96 58.7462.41Table 4. ANOM for the Means.259S/N 1 2 3Inner Gas Laser pressure Power Watt 36.42 40.08PSI 38.9539.36 -40.96Feedrate Powder Outer gasSpotOverlapFlow rate pressure Diameter FactorIPM 38.11 39.6239.74g/min 36.12 39.6741.68PSI 38.66 39.0539.76mm 38.7239.70 39.04% 37.37 39.7540.34Table 5.

Factor effect Plots for the Signal to Noise Ratios.In the above figures (3 and 4), IG / OG are Inner Gas / Outer Gas (PSI), P is Laser Power (W), F is the Feedrate (IPM), PF is the Powder Flowrate (g/min), D is the Spot Diameter (mm), O is the Overlap (%) and INT denotes the Interactions and Noises (their effects and percentage contributions would be calculated for further evaluation).260Optimal LevelsFrom the plots and the Analysis of the Means, it was found that the optimal levels that are required to produce the given OEC 1 are IG2= 4 PSI, P3 = 1000W, F3 = 30 IPM, PF3 = 17.5 g/min, OG3 = 12 PSI, D2 = 0.74mm and O3 = 45%. With these optimal levels the expected performance should be predicted and the confirmation experiment should be carried out to verify the validity of the results.From OEC 2, the study of the ANOM shows that the optimal levels remain the same except for the Feed rate and the Overlap factor.

It must also be noted that there are only slight differences in the S/N ratio values of those control factors.Predictive ModelThe predictive model is constructed from the ANOM results and is formed by considering the optimum level contribution of each factor to the deviation from the overall mean value for the experiment. The general formula for a predictive value is given below: [6]y A B C D ( ,,,)(cid:32)y(cid:14)(yA(cid:16)y)(cid:14)(yB(cid:16)y)(cid:14)(y C(cid:16)y)(cid:14) (4)...Where, y is a Quality Characteristic and A, B, C, etc.

are the control factors.In this experiment the predicted value for the S/N ratio was calculated to be 45.6 and the results of the confirmation experiment showed an S/N ratio of 45.2 which showed that the experiment was valid.Analysis of Variance (ANOVA)ANOVA is a statistically based decision making tool used for detecting any discrepancies in the average performances of the groups of data tested. ANOVA breaks down total variation into comprehendible sources. In other words it enables us to quantitatively estimate the relative contribution each control factor makes to the overall performance. This contribution is expressed as a percentage. The importance of each control factor is measured by comparing the variance between the control factor effects with that of the experimental data.

This contribution is expressed as a percentage. The importance of each control factor is measured by comparing the variance between the control factor effects with that of the experimental data. Here a mathematical technique known as the Sum of Squares (SS) is used to measure the deviation of the control factor effects on the average response quantitatively, from the overall experimental mean response. The effects of random experimental error can also be determined in this process. The basic steps involved in ANOVA are firstly determining the Grand Total Sum of Squares (GTSS).

The effects of random experimental error can also be determined in this process. The basic steps involved in ANOVA are firstly determining the Grand Total Sum of Squares (GTSS). The GTSS is comprised of the overall experimental mean (or) the Sum of Squares due to the mean which is given by [6]: NGTSS(cid:32) (cid:166)i1 (cid:32)(S N /)2 i(5)And the Sum of Squares due to the variation about the mean (or) the total Sum ofSquares is given by [6]:NTotal SS(cid:32)(cid:166)i1(cid:32)(S N /i(cid:16)S N /)2(6)It is important to note that GTSS = Total SS + SS due to the mean (See table 6).261Grand Total Sum of Squares Total Sum of Squares(dB)2 27845.98 250.27Sum of Squares due to the Mean 27595.71Table 6.

The Sum of Square values obtained from the experiment.The Sum of Squares due to variation about the mean for a Factor A can be given as [6]:(cid:167) (cid:168) (cid:168) (cid:168) (cid:168) (cid:169)(Number of experiments at ALevel 1SS = (Number of experiments at AALevel 2...(Number of experiments at A(cid:14)) *(S N /Level 1)*(S N /Level 2Level N)*(S N /(cid:16)S N /(cid:16)S N /Level N(cid:16))2(cid:14))2(cid:14)S N /)2(cid:183) (cid:184) (cid:184) (cid:184) (cid:184) (cid:185)(7)Percentage Contribution of ParametersThe percentage contribution is the portion of the total variation that was observed in an experiment and is attributed to each significant factor and/or interaction. It is a function of the sums of squares of each factor. If the factor and/or interaction levels were controlled precisely, then the total variation would reduce by the amount indicated by the percentage contribution.

If the factor and/or interaction levels were controlled precisely, then the total variation would reduce by the amount indicated by the percentage contribution. The percentage contribution due to error provides an estimate of the adequacy of the experiment. It is given by [6]:Percentage Contribution(cid:32)(SSfactor/Total SS) *100(8)Percentage SS for Factors Contribution SSIG SSP SSF SSPF SSOG SSD SSO SSINT SSError0.77 69.52 9.92 95.00 3.75 3.02 29.81 12.420.31 27.78 3.96 37.96 1.50 1.21 11.91 4.9626.0810.42Table 7. Sum of Squares due to each factor and their Percentage Contribution.From the Percentage Contributions (table 7 and figure 5) we may note that there are more unaccountable errors contributing around 10.4% to the overall variation along with the 4.96% of assumed interactions and uncontrolled noises. This is attributed to three main causes:(cid:120) Uncontrollable noise factors.

This is attributed to three main causes:(cid:120) Uncontrollable noise factors. (cid:120) Factors which are not included in the experiment and (cid:120) Experimental error [5] [7].262Percentage Contribution of Factors10%0%5%28%Inner GasPowerFeedrate12%Powder FlowrateOuter Gas1%Spot Diameter1%4%OverlapINTError39%Figure 5. Pie chart showing the Percentage Contributions.Number of Experiments = 18DOF due to IG = 1DOF due to D = 2DOF due to the mean = 1DOF due to INT = 2Total DOF = 17DOF due to O = 2DOF due to Error = 2DOF due to P = 2DOF due to OG = 2DOF due to PF = 2DOF due to F = 2Figure 6. Schematic of the Degree of Freedom decomposition [6].From figure 6, it is found that the DOFs due to all the factors taken into consideration are two less than the total DOF. This means that there are experimental errors and unknown or uncontrollable factor contributions worth two DOFs [7].

This means that there are experimental errors and unknown or uncontrollable factor contributions worth two DOFs [7]. In this situation, it is necessary to do an F-Test to validate the importance of the control factors and prove that they can still control the process in the presence of these errors and variations.263F- TestThe F-Test or F-Ratio, also known as the variance ratio, is used to test thesignificance of the factor effects. It is given by [6]:F(cid:32)MS 2 S e(cid:32)mean square due to a control factor mean square due to experimental error(9)Where,MS(cid:32)factor effect sum of squares factor degrees of freedom(9.1) &S2 e(cid:32)error sum of squares error degrees of freedom(9.2)When the value of F is more than 1, then the effect of the control factor is more than the variance due to experimental error and the interaction effects.

Some general guidelines for the F- Ratio are [6]:(cid:120) F<1, then the experimental error outweighs the effect of the control factors. The controlfactors will be trivial and indistinguishable from the experimental error.(cid:120) F (cid:124) 2, the control factor only has a reasonable effect compared to the experimental error. (cid:120) F>4, the control factor is strong compared to the experimental error and is obviouslysignificant.It is recommended to pool the insignificant control factors with the error. This makes it possible to calculate the contribution of the significant control factors more effectively [6] [7]. In this experiment, laser power, powder flow rate and overlap will be counted as significant factors and the remaining factors will be pooled together as the error as shown in table 8.Factor Mean Square2 Se5.08609MSP MSPF MSO34.7592547.4982414.90408Table 8.

Mean Square values for the significant factors and the error.The F ratios that are calculated from the above values (FP = 6.8, FPF = 9.3& FO = 2.9) show that Laser Power and Powder Flowrate are clearly significant when compared to the errors and Overlap factor has only a reasonably significant effect on the errors.ConclusionThe LAMP process at UMR was optimized using the Design of Experiments approach based on Taguchi’s methods for maximum build rate, hardness and minimum porosity within the level settings that were decided for this experiment. The control factor interactions were studied and it was shown that the contribution of some control factors namely the laser power, the overlap factor and the powder flowrate are significant and the optimal levels for this experiment were determined.

The contribution of the errors which include the experimental error, error due to interactions between factors and error due to uncontrollable noises were also studied and the system was made robust to these variations. A comparative study was also done with equal weightage for all the Quality Characteristics and the optimal values were not found to vary by much.264It is required that this process be further improved. This can be done by using the results of this experiment and conducting iterative experiments to hone in on the best possible level values of the control parameters that are least sensitive to the variations and are more improved and repeatable.References1. Liou, F. W., Choi, J., Landers. R. G., et al., “Research and Development of a Hybrid Rapid Manufacturing Process”, Twelfth Annual Solid Freeform Fabrication Symposium, Austin, Texas, August 6–8, 2001, pp. 138–145.2. Duley W. W., “Laser Processing and Analysis of Materials”, Plenum Press, New York, 1983, pp.

Creveling, “Engineering Methods for Robust Product Design Using Taguchi Methods in Technology and Product Development”, Addison Wesley Publishing Company, Massachussets, 1998.7. Ross, R. J. “Taguchi Techniques for Quality Engineering”, McGraw-Hill, New York, 1989.AcknowledgmentsThis research was supported by the National Science Foundation Grant Number DMI-9871185, Missouri Research Board, and Army Research Office. Their support is gratefully appreciated. The authors would also like to thank Gail Richards for proofreading the paper and enhancing its overall quality.265.

of Mech. and Nuclear Engineering, The Pennsylvania State Univ., Univ. Park, PA 16802 4 Applied Research Laboratory, The Pennsylvania State Univ., Univ. Park, PA 168021 AbstractThis work applies statistical analysis and uncertainty quantification tools developed for characterizing virtual microstructures in three dimensions to a two-dimensional experimental investigation of Ti-6Al-4V Widmanstätten ⍺-lath thicknesses obtained from back-scattered electron (BSE) or electron back-scatter diffraction (EBSD) images on two thin-walled samples manufactured via the LENS® process. The Materials Image Processing and Automated Reconstruction (MIPAR™) software optimizes unique recipes for conversion of the BSE or EBSD images to binary data, and subsequently computes the inverse of the linear intercept for each ⍺-lath.

Mean ⍺-lath thicknesses and discrete probability density functions (PDFs) of inverse intercepts are used to make quantitative comparisons of ⍺-lath structures at different heights throughout the thin walls. Real-time thermal data collected during the LENS® experiment is then compared to quantitative microstructural results in order to determine trends between ⍺-lath structures, thermal gradients, and melt pool areas across experimental process parameters.2 IntroductionDirect metal additive manufacturing (AM) refers to any process that manufactures near-net shape components via direct deposition of metallic material into a pool of molten metal. These processes differ primarily in the power source (laser- or electron-beam) and feedstock delivery system (powder stream, wire feed, powder bed). For each type of process, microstructural characterization techniques are used to assess and qualify the strength, reliability, and overall quality of metallic components.

For each type of process, microstructural characterization techniques are used to assess and qualify the strength, reliability, and overall quality of metallic components. Further, while AM has advantages over traditional machining processes, such as lead time improvements and lower per-unit costs for relatively small batches, the inherent temperature variability in all cyclic direct metal AM processes often results in inconsistent and undesirable solidification microstructures.One approach for understanding temperature- and solidification microstructure-variability indirect metal AM is a method termed “process mapping” developed by Beuth et al. [1]. Thismethodology maps process outcomes in terms of process variables, where variables represent experimentally-controlled factors such as beam power and velocity and outcomes refer to experimental results such as stress or microstructural characteristics.

The idea behind the process mapping approach is to allow all direct metal AM processes, which act in different regions of processing space, to be characterized and analyzed in a cohesive way. Early work focused on residual stress in metal and polymer AM processes [2-6]. Subsequent work investigated control of melt pool dimensions under steady state conditions for the full range of the Laser Engineering Net Shaping (LENS®) process through investigation of nondimensional thermal metrics [7-9]. This work was extended to provide direct predictions of solidification microstructure [10,11]. Additionally, many process mapping approaches have leveraged analytical and thermal finite element models to predict cooling rates and thermal gradients, parameters that are used to predict trends in solidification microstructures [12-16].

Further, within recent years, detailed microstructural observations have shown correspondence to theoretically calculated melt pool geometries for given sets of process variables [17-18].Real-time sensing and adaptive control utilizes process mapping techniques by monitoring deposition and altering process variables in order to maintain constant melt-pool geometry or alternatively to control component build parameters. Over the years, significant research regarding adaptive control of laser processes has been conducted in which monitoring methods are based on the physical phenomena that occur due to laser-materials interaction (see Ref. [19] for a recent thorough review). Many of the techniques and sensors used for investigation of AM processes have historically been used for laser cladding and welding processes (i.e. optical and acoustic techniques).

Many of the techniques and sensors used for investigation of AM processes have historically been used for laser cladding and welding processes (i.e. optical and acoustic techniques). For laser AM processes, researchers have leveraged optical imaging techniques to monitor temperature during laser-based manufacturing of various materials [20]. Griffith et al. studied melt-pool temperatures and gradients collected during deposition of H13 tool steel [21], and Hofmeister et al. achieved significant stabilization of melt-pool size via control of laser power in 316 stainless steel through use of an optical thermal imaging system [22]. Correlation of thermal data to solidification microstructure has also been a topic of interest for various materials, including 316 stainless steel [23] and Ti-6Al-4V [24-26, 27-28].

Correlation of thermal data to solidification microstructure has also been a topic of interest for various materials, including 316 stainless steel [23] and Ti-6Al-4V [24-26, 27-28]. However, most microscale observation of AM components is qualitative in nature and does not lend itself to the process mapping approach.In contrast, this work seeks to provide a methodology for the quantitative analysis of AM Ti-6Al-4V sub-grain level microstructures, using two thin-walled L-shaped components built via the LENS® process, and leveraging techniques similar to those utilized for quantifying the average size of Widmanstätten ⍺-laths in forged billets of Ti-6Al-4V [29]. The Materials Image Processing and Automated Reconstruction (MIPAR™) [30] software was used to optimize recipes for conversion of back-scattered electron (BSE) and electron back-scatter diffraction (EBSD) images into binary data suitable for mean linear intercept (MLI) [31] analysis and thus ⍺-lath thickness calculations.

In addition, a quantitative comparison technique for ⍺-laths observed in different regions of the two components is presented, where discrete PDFs of MLIs are compared for geometric similarity via a Modified version of the Bhattacharyya Coefficient(MBC) [32]. Lastly, quantitative ⍺-lath data is analyzed for trends relative to the real-time thermal data collected during the 3D directed energy deposition experiment used to create the two thin walls. Kriczky et al. published an overview of the experiment [33] and this work compares thermal gradients and melt-pool areas to ⍺-lath thicknesses in the resulting solidification microstructure for both thin-walled L-shaped components.3 Additive Manufacturing and Thermal Imaging of Ti-6Al-4V ComponentsThe LENS™ system used to build components investigated in this work represents a collaborative effort between Penn State University’s Center for Innovative Metal Processing through Direct Digital Deposition (CIMP-3D), Stratonics, and Optomec.

The integrated multi- sensor/process model-based control system leverages an Optomec MR-7 LENS™ system, outfitted with a Stratonics ThermaViz® thermal imaging camera. Using the integrated system, CIMP-3D manufactured the two relatively-large (~50 mm tall), L-shaped, thin-wall samples of AM Ti-6Al-4V, shown in Figure 1. Each sample is composed of two legs; the first leg is constructed from a single bead, while the second has three beads per layer. The key difference between these samples is the dwell time used between added material layers. One sample was built with zero dwell time between layers, while the other sample had a dwell time of four seconds. These two samples are shown in Figure 1A and Figure 1B, respectively.Figure 1: AM Ti6Al-4V components constructed using the LENS™ process: L-shaped thin- walled samples under investigation; 0 s dwell (A), and 4 s dwell (B) [33].The complete thermal history of each component was captured in real time during theexperiment.

Detailed information regarding the step-by-step procedures to transform point-by- point thermal data into thermal gradients, melt pool areas and the like is available in [33] for the interested reader. Here, only the authors’ designations for the lower 10%, upper 90%, and center 33% are of interest for both the 1- and 3-bead legs of each sample; these designations are illustrated using the zero second dwell specimen shown in Figure 2.

For this work, ⍺-lath datawas collected for both legs of each sample (0s dwell, 4s dwell) within the center 33% region and along the entire height of the thin walls (z-direction as pictured below).Figure 2: Thermal gradient reconstruction showing spatial designations for AM Ti-6Al-4V components: Isometric 3D reconstruction of calculated thermal gradients for the zero second dwell sample [33].4 Microstructural Characterization of Ti-6Al-4V ⍺-lathsIn alpha-beta titanium alloys such as Ti-6Al-4V, the body-centered cubic (BCC) β (beta) phase transforms to a hexagonally close packed (HCP) ⍺ (alpha) phase upon cooling through the β transus, which occurs at approximately 1270° K [35]. The smaller ⍺ features of Ti-6Al-4V begin growing within the metastable β phase from grain boundaries to produce ⍺ lamellae/laths within a β matrix. The overall ⍺+β morphology that results is called the Widmanstätten morphology, which can take two forms: colony or basketweave ⍺.

The overall ⍺+β morphology that results is called the Widmanstätten morphology, which can take two forms: colony or basketweave ⍺. In AM processes, depending on part geometry and AM process parameters, cooling rates can vary significantly across a single part, especially in between added layers and on the edges of parts. As a result, multiple ⍺ structures are often observed in single β grains, as exemplified in Figure 3 which shows two BSE images depicting colony- and basketweave-⍺ in the same field of view.

As a result, multiple ⍺ structures are often observed in single β grains, as exemplified in Figure 3 which shows two BSE images depicting colony- and basketweave-⍺ in the same field of view. While microstructural variation is characteristic of the LENS™ samples under investigation, such inconsistent microstructure does not lend itself to rapid qualification methods for AM nor does it make quantifying microstructural features using existing toolsets a straightforward task.Figure 3: Colony and basketweave ⍺ microstructures observed in LENS™ Ti-6Al-4V: BSE image showing colony ⍺ microstructures present amidst a majority of basketweave ⍺; observed in the 1-bead leg of the sample with zero dwell time between layers.Inspection of both samples and their solidification microstructures resulted in a designationof five zones along the height of each sample, as pictured in Figure 4 below. Additionally, this inspection yielded the best way to image ⍺-laths of various sizes.

Additionally, this inspection yielded the best way to image ⍺-laths of various sizes. For all sizes, very fine polishing and thorough cleaning was required, however the mode of data collection changed for different lath sizes. For thicker laths like those pictured in Figure 3, back-scattered electron (BSE) imaging at a magnification of 2000x (working distance 5 mm, accelerating voltage 20 kV) proved to be a repeatable methodology. However, smaller laths required crystallographic analysis via electron back-scatter diffraction (EBSD) (step size 0.1 μm) to be distinguishable.Figure 4: Zones 1-5 used for ⍺-lath width analysis: Spatial locations along the sample height for zone designations used in ⍺-lath width analysis.Zones 2, 3, and 4, those not located at the extreme ends of the sample (top, bottom), appeared to have steady state microstructures.

More specifically, the central zones along theheight of the thin walls appeared to have qualitatively similar lath sizes, which can be seen in Figure 5. However, zones 1 and 5 were very difficult to image using BSE techniques, and produced BSE images unsuitable for image processing. As a result EBSD data was required, as shown in Figure 6.

However, zones 1 and 5 were very difficult to image using BSE techniques, and produced BSE images unsuitable for image processing. As a result EBSD data was required, as shown in Figure 6. The laths requiring EBSD appear smaller based on the more acicular nature of the observed ⍺-microstructure.Figure 5: Steady state zones of AM Ti-6Al-4V components: BSE images of zones 2, 3, and 4 showing easily quantifiable basketweave ⍺ structures; taken from the 1-bead leg of the sample with zero dwell time between layers.Figure 6: Top and Bottom Zones of AM Ti-6Al-4V components, EBSD Imaging: EBSD images of zones 1 and 5 showing more acicular microstructures than those in other zones; taken from the 1-bead leg of the sample with zero dwell time between layers.5 Image Processing via MIPAR™Lath thicknesses and distributions are determined using the image analysis software MIPAR™ (Materials Image Processing and Automated Reconstruction) [30], developed by Sosa, et al.

at The Ohio State University. For each imaging technique, a series of imageprocessing steps was completed within MIPAR™ to convert the data into binary images for quantification.For BSE imaging in steady state zones, a 6-step series of image filtering techniques was implemented to convert grayscale images to binary, as shown in Figure 7. In particular the following series of filters was used to segment features: Wiener filter, Fast Fourier Transform filter, adaptive threshold, dilation, erosion, and feature rejection.

In particular the following series of filters was used to segment features: Wiener filter, Fast Fourier Transform filter, adaptive threshold, dilation, erosion, and feature rejection. Although filter choices were based on experience and trial-and-error, the optimum parameters for each filter were objectively determined based on segmentation-to-image mutual information for each image-processing algorithm used in the sequence (see Ref [30] for details).Figure 7: BSE image filtering for steady state zones: Example of a BSE image converted to binary for ⍺-lath width analysis using MIPAR™.Colored EBSD images were converted to grayscale from their default RGB format by adding(R+G+B) content for each pixel. Similarly to the images in the steady state zones, a unique multi-step image-processing sequence was developed.

Similarly to the images in the steady state zones, a unique multi-step image-processing sequence was developed. Specifically, a median filter and edge find sequence was applied for three rounds, followed by a final erosion and feature rejection, producing binary data as shown in Figure 8. Optimum filter parameters were again determined based on segmentation-to-image mutual information.Figure 8: EBSD image filtering for top and bottom zones: Example of an EBSD image converted to binary for ⍺-lath width analysis using MIPAR™.6 Determining ⍺-lath Thicknesses and Quantifying Uncertainty in MicrostructureHaving obtained binary images from the microstructural data, Widmanstätten ⍺-lath thicknesses were calculated via the mean inverse of the linear intercept [31], a procedure also performed within MIPAR™.

This method superimposes a grid of parallel lines on the image, subsequently rotates the grid by 10° increments to collect measurements in multiple directions, and records linear intercepts (λ) and inverse intercepts (1/λ) computed for each lath (see [29,31]) so that the true 3D thickness of ⍺-laths can be estimated via Eq. (1).𝑡ℎ𝑖𝑐𝑘𝑛𝑒𝑠𝑠 =1 1.5(1/𝜆)𝑚𝑒𝑎𝑛(1)The application of this method approximates laths as layered structures and estimates the truethickness of a series of “infinite” plates of finite thicknesses (i.e. finite only in 1 dimension). Further, it works best for more randomly oriented (i.e. 100% basketweave) structures; thus the ⍺ colonies visible in Figure 3 provide an additional source of uncertainty in these measurements.Average lath thicknesses, computed for each image and averaged based on the number ofimages taken for each section are shown in Table 1 below.

The confidence intervals were computed using Student’s t-distribution on (n-1) degrees of freedom as a benchmark with a significance level of 0.05 to produce two-sided 95% confidence bounds.

Inverse intercept distributions are constructed using all inverse intercepts calculated for the set of images collected in each zone (see Table 1 above for the number of images collected in each set). Discrete distributions are binned using a 30-bin scheme between 0 and 5 that was based on the observed data, which typically ranged between 0 and 5. The larger-valued outliers (corresponding to extremely small alpha lath widths) inherently result from this analysis due to complicated three-dimensional basketweave microstructure being observed in only two dimensions, although imperfect binary representations likely play a role as well. Once constructed, the distributions within the zones of each sample leg are compared for geometric similarity via a Modified version of the Bhattacharyya Coefficient (MBC) [32], which is applied to discrete PDFs via Eq. (2).𝑛 𝑖=1𝑀𝐵𝐶 = √1 − ∑ √𝑅𝑖𝑆𝑖(2)In Eq.

(2).𝑛 𝑖=1𝑀𝐵𝐶 = √1 − ∑ √𝑅𝑖𝑆𝑖(2)In Eq. (2) Ri and Si correspond to the fraction of inverse intercept data contained in bin ifor the discrete probability distributions (i.e., histograms) R and S, and correspond to the two distributions being compared. For this work, distributions are compared within sample-leg steady state zones (i.e. zone 2 to 3, zone 2 to 4, and zone 3 to 4) as well as between sample-leg top and bottom zones (i.e. zone 1 to 5). Comparing BSE and EBSD results separately ensure that no aliasing effects occur based solely on variations in data collection procedures and image processing techniques.Although other methods of discrete distribution comparisons exist, the MBC was selected for its relatively high sensitivity to subtle variations in the PDFs. Additionally the MBC statistic ranges between 0 and 1, where a value of 0 implies that two models are identically distributed, and higher values imply greater variance between distributions.

Additionally the MBC statistic ranges between 0 and 1, where a value of 0 implies that two models are identically distributed, and higher values imply greater variance between distributions. For reference, previous work by this author has shown that an MBC≤0.13 produces qualitatively well-matched distributions [34].For each leg, MBC calculations between zones are shown in Table 2. MBC values abovethe threshold of 0.13 are highlighted in red, and in each of these cases a clear qualitative difference in microstructure can be observed in BSE or EBSD images. The highest MBC in Table 2, corresponding to a comparison of zone 1 vs.

The highest MBC in Table 2, corresponding to a comparison of zone 1 vs. 5 in the 1-bead leg of the 4s dwell sample, was determined to be attributable to poor data collection quality of the EBSD scan used for analysis of zone 1.Table 2: MBC comparisons of inverse intercept distributions: MBC calculations for each sample leg, comparing top and bottom regions as well as steady state zones.Zone 1 - Zone 5 Zone 2 - Zone 3 Zone 2 - Zone 4 Zone 3 - Zone 41-Bead Leg, 0s Dwell 0.045 0.195 0.189 0.0291-Bead Leg, 4s Dwell 0.328 0.072 0.092 0.1523-Bead Leg, 0s Dwell 0.097 0.045 0.043 0.0753-Bead Leg, 4s Dwell 0.092 0.041 0.074 0.103Distributions of inverse intercept length can also be analyzed graphically for each leg.For brevity, only the 1-bead leg zero second dwell case is shown in Figs. 9-10. Zones 3 and 4, as well as zones 1 and 5 appear to have significantly similarities, respectively.

9-10. Zones 3 and 4, as well as zones 1 and 5 appear to have significantly similarities, respectively. These conclusions are supported by the data in Table 2 above, where the comparison of zones 3 and 4 produces an MBC = 0.029 and zone 1 to zone 5 produces and MBC = 0.045, both providing values well below the threshold of MBC≤0.13. Notably, these zones had extremely similar lath thicknesses, 0.69 vs. 0.67 µm for zone 3 vs. 4 and 0.55 vs. 0.59 µm for zone 1 vs. 5. Further parallels can bedrawn with the lath thickness data in Table 1. In particular, the distributions with higher fractions at smaller inverse intercepts are those with larger lath thicknesses.Interestingly, the inverse intercept lengths observed in particular bins of zones 1 and 5 foreach leg produced distributions that were not smooth in certain ranges (see Figure 10).

This occurred for all zone 1 and 5 data, and is perhaps due to the image processing routines used for analysis; however this is a bias left for investigation in future work.Inverse Intercept Distributions, 0s Dwell, 1 Bead Leg, Zones 2-4s t p e c r e t n I0.10e s r e v n IZone 2f o0.05Zone 3n o i t c a r FZone 40.000.001.002.00 Inverse Intercept Length (1/λ)3.004.00Figure 9: Inverse intercept distributions, zones 2-4, 0s dwell, 1-bead leg: 1/λ distributions plotted between 0 and 4.Inverse Intercept Distributions, 0s Dwell, 1 Bead Leg, Zones 1 & 5s t p e c r e t n I0.10e s r e v n If o0.05n o i t c a r F0.000.001.002.00 Inverse Intercept Length (1/λ)3.00Figure 10: Inverse intercept distributions, zone 1 vs.

5, 0s dwell, 1-bead leg: 1/λ distributions plotted between 0 and 4.7 Comparison of Widmanstätten ⍺-lath Widths to Thermal Imaging DataLastly, a comparison of the quantitative ⍺-lath thicknesses to the thermal imaging dataprovided in Ref [33] is in order. For these comparisons, zones 2-5 are grouped to correspond to the upper 90% region of the thermal imaging data, leaving zone 1 to represent the lower 10%. All microstructural analysis was completed within in the center 33% of each leg of the samples (see Figure 2 for reference), so Region 1 is designated as the lower 10% and center 33% area of the samples (i.e. zone 1) while Region 2 is the upper 90% and center 33% of the samples (i.e.

Consideration of the dimensionless thermal gradient and dimensionless cooling rate used by Bontha et al. [12] reveals that the thermal gradient in the horizontal (layer) direction scales with cooling rate. Recent unpublished theoretical work has also shown that the trailing edge thermal gradient at the solidus-liquidus interface occurs strictly in the horizontal (layer) direction at the top surface of the melt pool. Thus, this particular thermal gradient data, which was collected and from a downward-facing thermal imaging camera and computed at the top of the melt pool, can be interpreted in the same way as a cooling rate for trend identification purposes since the velocity of the beam used in this experiment was constant.

Therefore, ⍺-lath thicknesses appear to be larger when cooling rates are lower and smaller when cooling rates are higher, an intuitive result.In each of the tables above, thermal gradient increases as the melt pool area decreases.Noting Gockel & Beuth showed that lines of constant melt pool area correspond to lines of constant β grain size for wire-feed AM Ti-6Al-4V [1], it appears that larger melt pool areas correspond not only to enlarged β grain sizes but also to thicker ⍺-laths within those grains. Inspection of Table 4 and Table 5 reveals that ⍺-lath thicknesses increase with the number of beads used per layer and decrease with pause length.

Inspection of Table 4 and Table 5 reveals that ⍺-lath thicknesses increase with the number of beads used per layer and decrease with pause length. Both an increase in number of beads and a decrease in pause length correspond to an increase in the heat available in the system, providing a mechanism for continued ⍺-lath growth.It should be noted, however, that though these trends exist it can also be observed frominspection of the data that neither thermal gradient or melt pool area are sufficient to predict lath thicknesses. For example, the measured thermal gradients from the three-bead leg of the zero second dwell sample increased from region 1 to region 2; a result that is inconsistent with the other three sample legs. Another anomalous comparison can be observed from the three-bead leg of the four second dwell sample, where thermal gradients of 6925 and 6893 K/cm were observed in region1 and region 2, respectively, while alpha lath widths were computed to be 0.41 and 0.67 μm.

This suggests that more work is necessary to properly correlate in situ thermal measurements to ⍺ microstructures in AM processes, perhaps by considering thermal history below the β transus when lath growth is expected to be active.8 ConclusionsThis study presents a general framework for quantitative analysis of sub-beta-grain level Widmanstätten ⍺ microstructures in AM Ti-6Al-4V. The methodology includes techniques for analyzing ⍺-laths of various sizes, as well as an uncertainty quantification tool for comparing inverse linear intercept distributions in various regions of AM components. The quantitative results agree with qualitative visual analysis, and favorably correspond to published theoretical and experimental results. In confirming intuitive trends between thermal gradient, melt poolsize, and solidification microstructure, this work provides a quantitative means for verifying intuition about ⍺-level structure in AM Ti-6Al-4V.

However, although trends were identified, it was also observed that the neither thermal gradient or melt pool size alone is sufficient to predict ⍺-lath thicknesses, and thus additional research is required to further understand the complex thermal cycles present in AM-fabricated structures.9 AcknowlegementsThis material is based in part on research sponsored by Air Force Research Laboratory through America Makes under agreement number FA8650-12-2-7230. The U.S. Government is authorized to reproduce and distribute reprints for Governmental purposes notwithstanding any copyright notation thereon. The views and conclusions contained herein are those of the authors and should not be interpreted as necessarily representing the official policies or endorsements, either expressed or implied, of Air Force Research Laboratory or the U.S.

Government.10 References1Beuth, J., Fox, J., Gockel, J., Montgomery, C., Yang, R., Qiao, H., Soylemez, E., Reeseewatt, P., Anvari, A., Narra, S., and Klingbeil, N., “Process Mapping Across Multiple Direct Metal Additive Manufacturing Processes,” in Solid Freeform Fabrication Proceedings, Austin, 2013, pp. 655-665.2Klingbeil, N. W., Zinn, J. W., and Beuth, J. L., "Measurement of residual stresses in parts created by shape deposition manufacturing," in Solid Freeform Fabrication Proceedings, Austin, 1997, pp. 125-132.3Klingbeil, N. W., Beuth, J. L, Chin, R. K., and Amon, C. H., "Measurement and modeling of residual stress-induced warping in direct metal deposition processes," in Solid Freeform Fabrication Proceedings, Austin, 1998, pp. 367-374.4Ong, R., Beuth, J. L., and Weiss, L. E. "Residual stress control issues for thermal deposition of polymers in SFF processes," in Solid Freeform Fabrication Proceedings, Austin, 2000, pp. 209-218.5Klingbeil, N. W., Beuth, J.

432- 440.8Vasinonta, A., Beuth, J. L., and. Griffith, M., "A Process Map for Consistent BuildConditions in the Solid Freeform Fabrication of Thin-Walled Structures," ASME Journal of Manufacturing Science and Engineering, Vol. 123, pp. 615-622, 2001.9Vasinonta, A., Beuth, J. L., and. Griffith, M., "Process Maps for Laser Deposition of Thin-Walled Structures," in Solid Freeform Fabrication Proceedings, Austin, 1999, pp. 383-391.10Klingbeil, N. W., Brown, C., Bontha, S., Kobryn, P., and Fraser, H., "Prediction of Microstructure in Laser Deposition of Titanium Alloys," in Solid Freeform Fabrication Proceedings, Austin, 2002, pp. 142-149.11Klingbeil, N. W., Bontha, S., Brown, C. J., Gaddam, D. R., Kobryn, P. A., Fraser, H. L., Sears, J. W., "Effects of Process Variables and Size Scale on Solidification Microstructure in Laser-Based Solid Freeform Fabrication of Ti-6Al-4V," in Solid Freeform Fabrication Proceedings, Austin, 2004, pp. 92-103.12Bontha, S.

92-103.12Bontha, S. and Klingbeil, N.W., "Thermal Process Maps for Controlling Microstructure in Laser-Based Solid Freeform Fabrication," Solid Freeform Fabrication Proceedings, pp. 219- 226, 2003.13Bontha, S., Klingbeil, N.W, Kobryn, P., and Fraser, H.L., "Effects of Process Variables and Size Scale on Solidification Microstructure in Beam-Based Fabrication of Bulky 3D Structures," Materials Science and Engineering A, vol. 513-514, 2009, pp. 311-318.14Bontha, S., Klingbeil, N.W, Kobryn, P., and Fraser, H.L., "Thermal Process Maps for Predicting Solidification Microstructure in Laser Fabrication of Thin Wall Structures," Journal of Materials Processing Technology, vol. 178, 2006, pp. 135-142.15Davis, J., Klingbeil, N., and Bontha, S., "Effect of Free-Edges on Melt Pool Geometry and Solidification Microstructure in Beam-Based Fabrication of Bulky 3-D Structures," in Solid Freeform Fabrication Proceedings, Austin, 2010, pp.

230-241.16Davis, J., Klingbeil, N., and Bontha, S., "Effect of Free-Edges on Melt Pool Geometry and Solidification in Beam-Based Fabrication of Thin-Wall Structures," in Solid Freeform Fabrication Proceedings, Austin, 2009, pp. 320-331.17Beuth, J. L., "Methods for Determining Process Variable Combinations Yielding Constant Melt Pool Geometry for Single Bead Deposits in Direct Digital Manufacturing," Materials Science and Technology, Vol. 31, Issue 8, June 2015, pp. 912-916.18Gockel, J., and Beuth, J., "Understanding Ti-6Al-4V Microstructure Control in Additive Manufacturing via Process Maps," in Solid Freeform Fabrication Proceedings, Austin, 2013, pp. 666-674.19Purtonen, T., Kalliosaari, A., Salminen, A., “Monitoring and adaptive control of laser processes,” Physics Procedia, Volume 56, 2014, pp.

M., Kampe, S. L., Microstructural evolution in laser-deposited multilayer Ti-6Al- 4V builds: Part I. Microstructural characterization,” Metallurgical and Materials Transactions A, Vol. 35, Issue 6, June 2004, pp. 1861-1867.Kelly, S. M., Kampe, S. L., Microstructural evolution in laser-deposited multilayer Ti-6Al- 4V builds: Part II. Thermal Modeling,” Metallurgical and Materials Transactions A, Vol. 35, Issue 6, June 2004, pp. 1869-1879.Collins, P. C., Welk, B., Searles, T., Tiley, J., Russ, J. C., Fraser, H. L., “Development of methods for the quantification of microstructurl features in ⍺+β-processed ⍺/β titanium alloys,” Materials Science and Engineering A, Issue 508, 2009, pp. 174-182.Sosa, J. M., Huber, D. E., Welk, B., and Fraser, H. L., “Development and application of MIPAR: a novel software package for two- and three-dimensional microstructural characterization,” Integrating Materials and Manufacturing Innovation, Vol. 3, Issue 10, April 2014.Gundersen, H. J.

e Application of an Artificial Body Force to the Selective Laser Sintering Processby Lawrence S. Melvin TIl and Dr. J. J. Beaman University of Texas at Austin Department of Mechanical EngineeringAbstractAn artificial body force generated by a magneticfield is applied to the green powder bed of a ferromagnetic powder during the Selective Laser Sintering process. Preliminary experiments and theory are formed to determine whether the artificial body force is beneficial to the Selective Laser Sintering process and if it is usable within that process. Several applications are discussed including microgravity situations and two phase sintering processes. It is determined that the magnetic body force is beneficial to the Selective Laser Sintering process.IntroductionThe purpose of this experiment is to determine if an artificially generated bodyforce applied to the powder bed during application and sintering will enhance the Selective Laser Sintering process.

The force which is currently under consideration is a magnetostatic force applied to ferromagnetic materials. The addition of the artificial body force is expected to increase the green bed and part densities and therefore increase the part strength of the sintered work piece. In addition, the magnetic force is expected to induce a packing structure into the green powder bed. This packing structure generation was exhibited in an electrostatic powder application system, and therefore is expected in a magneto static system. Also, the magnetostatic system has the ability to generate a constant and controllable force during the sintering process.The magnetostatic system is intended to be used to enhance part quality, replacegravity in a microgravity situation, and aid the powder application and leveling processes. The most important disadvantage of the magnetostatic force generation system is its nonlinearity with respect to radial position from the solenoid.

The most important disadvantage of the magnetostatic force generation system is its nonlinearity with respect to radial position from the solenoid. This nonlinearity will eventually have to be designed out of the system in order to produce high quality parts. If a satisfactory field cannot be generated, the use of a magnetic body force in the Selective Laser Sintering process will not be feasible.TheoryThe magnetostatic system is currently being modeled in simple terms. This model will illustrate qualitatively how the particles are expected to behave in the magnetic force field. This information can then be used to design an appropriate core for magnetostatic force field generation.The first step in generating a workable magnetic force field is to select the core parameters which are desired. The manipulation of the core parameters N, i, a, and d allow for the generation of a useful magnetic force field.

The manipulation of the core parameters N, i, a, and d allow for the generation of a useful magnetic force field. A standard model of the magnetic field intensity H for a single solenoid on the vertical axis is presented below:118Ni 2dz ad Z -+ a 2ad--;:=====- +1--;::======~ 1+(1a -iY ~ 1+(2: +iYThe magnetic core and sintering platform parameters are illustrated in Figure 1 below.zFigure 1 Solenoid calculation parameters N is the number of turns on the solenoidand iis the desired currentAs the value of ffa «1 the model can be reduced to a more simple form,Ni Hz: 2a(2).Equation (1) has been simulated to aid in the design of a suitable solenoid for the magnetic process. The parameters of N, a, d, and i were varied, and Hz as a function of z was generated and graphed. Many of the possible combinations of these parameters were simulated, but four general trends are illustrated here.

Many of the possible combinations of these parameters were simulated, but four general trends are illustrated here. A graph is presented below in Figure 2.Several important facts can be learned from this simulation. The first fact issimply that the parameters N and i only affect the magnitude of the field. However, this simulation also illustrates that a and d can be manipulated to effect where and how the field breaks from its initial linear behavior to fall to small values of field intensity. This phenomena allows a design to be implemented which will produce a constant linear force by operating in the region before the field breaks.119(1)Magnetic B Field vs.

: : ! ~..:.:.:..-:-Tu:=~:::·:·~·~r·~·:·:··:·:~·::u:.: '"I j : : ~ :1__I - - j - - : : ~ i1: : :0.008.000.020.040.060.080.10z (meters)Figure 2 The magnetic field intensity as a function of vertical displacement along the z axisThe force exerted on a ferromagnetic particle in this magnetic field can also be expressed mathematically. A simple model of the magnetic force is used to understand theses interactions.(3)The constant, c, is a function of the windings and geometry of the solenoid. This model can be coupled with the magnetic field intensity relationship to produce a force model involving the magnetic field intensity.F_ 4acHz2 z- N2z(4)This relationship provides an insight to how the particles are affected along the vertical axis. As the particles are distributed radially further from the axis, the force influencing them will lessen. This fact leads to a potential problem during the application of the particles.

As the particles are distributed radially further from the axis, the force influencing them will lessen. This fact leads to a potential problem during the application of the particles. It is possible and probable that this effect will lead to nonuniform powder layer distribution and possibly undesirable particle migration during the sintering process. This can be counteracted by designing an axially symmetric field and possibly by designing the Selective Laser Sintering process control software to account for this distortion. However, this problem is beyond the initial investigation which this experiment encompasses, and therefore it will be addressed in later work.120nA3"W....'".u·""'A'l.rI on a one kilowatt laser in the HighI1l"u",""",,'-W 1-0:,at Austin. A solenoid consisting of put into the part cylindercore was builtO·t"hVA3t"t to the top and bottom of theleV~~lea by hand to a thickness of 0.020 inches.noheat source.

A solenoid consisting of put into the part cylindercore was builtO·t"hVA3t"t to the top and bottom of theleV~~lea by hand to a thickness of 0.020 inches.noheat source. One layer was1" 1n1("111111 ITsolenoid...........,J"o>"...''''' .....'''....'-"' powder and the ............................h platform on thewhile the non magnetic along the line of sinter,triangles, while the..:P1J"'......... ,a ............:P were fragile and full of holesU .... IL4-l......,......p, ............# ........, v .... Microscope, more the two sinters. TheC1Tl·f'4001"".rI to one or two other grains while the four, or five different grains.LU, . . . 'L............, ..... tothan those of the3no magnetic field.OIn..... -w ...-woril 2000xFigure 4 Nickel powder sintered with 1.5 A of current producing a magnetic field. Magnified 2000xnonmagnetized sinter. In addition, there were many nickel particles in the magnetized sinter which. had numerous groves in them.

Magnified 2000xnonmagnetized sinter. In addition, there were many nickel particles in the magnetized sinter which. had numerous groves in them. However, most of the particles of the nonmagnetized sinter had smooth surfaces, and those which did not had only a small number of groves on them The magnetized sinter also exhibited some ridge like patterns in the sintered work piece.Experimental VariancesThree different experimental variances may have contributed to the observedexperimental results. First, the nonreflective coating was missing from the inside of the zinc solenide window during these experiments. Secondly, the powder was jostled about between the leveling and sintering processes when it was installed into the sintering machine. Finally, the solenoid generates heat which transfers to the sintering area through the sintering platform.The missing window coating has made the effective sintering power lower than itwas set on the machine.

For this reason, the balling effect which is exhibited when sintering single layers of nickel was not exhibited and it may be deduced that the entire region being sintered did not change to liquid phase during the heating process. Also, the exact power which reached the sintered surface cannot be determined in a reliable fashion. Because of these problems, these experiments are only comparable to122two variances cannotlinked to any effects on the sintered......' " " " J' " 'li1l:111-hI"\1111"rI',.n .."'1..::'UI!J•.:Jw,u""u. conclusions."... ........ ""h·..·.... r ..... that a magnetic field produces a Sintering process. The SEM malgnetl:i~ea sinter is more dense than thebonds are bigger, and there are morebonds 1"'I>T'J'l'C'131'lt the .........,jiO,"..'v .......,,""u. layer. influence conductthe to penetrate further into the magnetizedeffect and greater thicknessgaps due to ""11"........... 0+.. ""'....

These twowhich candesigned to meet specific material property----.---J and its potential in the powdergravitational bodyuse of the magnetic body force to during powder application and sinteringC"~ .... 'tClOT''''.ridifferent environments including low gravity and upside effect of the magnetic field will lead to greater part as well as enabling the production of parts which are..,""1I"C'1T·....11...............'v ....,u,..""u. parts.Hammond, NY,~~~~~~~~~~!;,2, The Macmillan Company, New York,Haus, tteJrID2U1Melcher, Electroma~netic Fields and Ener~y, PrenticeEnglewoodcliffs, NJ, 1989, pp. 318-9.~~~~~~~~~i:l., Prentice Hall, Engelwoodcliffs, NJ,Melvin,Electrostatic Application of Powder for Selective LaserUniversity ofat Austin, 1991..

plication of Cementitious Bulk Materials to Site ProcessedFreeform ConstructionJoseph Pegna Department of Mechanical Engineering, Aeronautical Engineering and Mechanics Rensselaer Polytechnic Institute Troy, New York 12180·3590 U.S.A. (518) 276·6030 Fax: (518) 276·2623 email: pegnaj@rpLeduAbstract: This paper reports a radical departure from generally accepted concepts in construc tion automation and demonstrates that new techniques of layered manufacturing can be applied effectively to construction. In the process, we also modified material processing of cement to adapt it to the requirements of Solid Freeform Fabrication.Our purpose is illustrated with sample structures manufactured by incremental deposition of reactive bulk materials (cement and Silica in this instance), a characterization of their material properties, and an assessment of their potential for Solid Freeform Fabrication of large struc tures.

For example, we estimated that it would take about months to build a structure the size of a 3000 sq. ft. house.Keywords: Selective aggregation, Large structures, Functional Prototypes, Multimodal struc tures, Construction Automation.1 Additive manufacturing processes are a recent addition to the engineer's arsenal. They result from a fundamental paradigm shift in manufacturing technology. Additive processes build-up a part in small organized increments rather than carving it out of a blank, as has been customary for centu ries. This new approach provides the designer not only with a tool for rapid prototyping and tool- opportunity to obtain functional parts and systems that are inconceivable by to obtain systems that are multimodal in terms of build mate- topologies, and that can customize material to the intended function of theMANUFACTURING PR'OCIES~)ESOFbut with means.

Ideally, one woulddesign.The fundamental principles involved in Additive Manufacturing Processes demand that solid material be added incrementally and eventually grown or assembled into the desired shape and material composition. The first dimension of this problem is geometric in nature depending upon scanning technique used for building up the part. The desired incremental geometry can be obtained by either sweeping a point, a line or a surface. The second dimension of the problem is transport to the deposition area. Material can be transported in gas or liquid form and selectively subjected to a change of phase; or it can be transported in solid phase and selectively1995 Joseph Pegna, printed with permission.39aggregated. All additive manufacturing processes can be characterized by their geometric and transport attributes as summarized in Tables 1 and 2.Our research program is concerned with the fundamental principles of additive manufacturing processes.

It seeks to achieve a process capable of providing not only geometry but also custom ize material properties within a generic environment. In view of the Rapid Prototyping/Solid Freeform Fabrication literature, it is possible to attain fine control over the nature and properties of the deposit obtained by selective change of phase (See [Maxwell et aI., 1995] for example.) We conjecture however that control of the material nature and microstructure limits admissible flow rates to about 108f..!m3Is. Beyond that rate, it appears that most researchers assume that additive manufacturing processes by selective change of phase are material specific. Hence specific pro cesses have been developed for specific materials (stereo1itography for example.)The most likely route to additive manufacturing for large parts -while retaining some mate rial flexibi1ity- is selective aggregation of powders.

Selective laser sintering, for example, has already been shown to apply to various polymer, ceramics and metal powders. The resulting structure however is a porous aggregate with inhomogeneous properties.2 MOTIVATION AND BACKGROUND Insofar, Solid Freeforrn Fabrication has proved useful in the rapid fabrication of visual prototypes and tooling. As the technology develops, it is now aiming at producing functional parts and sys tems. In this natural evolution of technology, parts are sti11limited in size to a few decimeters for non-functional prototypes and a few centimeters for functional ones. Can one process deliver functional prototypes of meter size and above? This is the challenge addressed by this project. It's eventual aim is to investigate the use of layered manufacturing as a potential approach to Con struction Automation.Indeed, construction may very well be the single most difficult challenge to manufacturing automation.

Construction is essentially a prototype fabrication, and has so far resisted most attempts at automation. Other manufacturing sectors have seen their productivity increase manyTABLE 1.Classification of Additive Manufacturing Processes by Selective Change of PhaseTABLE 2.Classification of Additive Manufacturing Processes by Selective AggregationPointI Ir;-IGas ....._--LCVD -SALD!LLCVD - Stereolithography~Ballistic Particle Manufacturing. -Sanders Prototyp- ing Inc.™ -Shape Melting.Fused Deposition Manufacturing.3-D Printing. - Selective Laser Sintering..~Photolytic LCVD of lines.NoneI~None.None.None.Blanket LCVDCubitaFM - Light Sculpting™ - Mask ElectroplatingSelective Aggre-~ J gation of Reactive i ~ ~Bulk Material (Presented in this paper)Filament Winding.Composite Pre- preg Fabrication. -Laminated Object Manufacturing. -Spray Deposition.40times over when they became automated.

-Laminated Object Manufacturing. -Spray Deposition.40times over when they became automated. The last 10 years of extensive experiments in construc tion robotics however, have yielded little or no productivity improvement [Kajioka et aI., 1990].The reasons for this apparent paradox is discussed in [Pegna, 1995]. In a nutshell, we claim that it is due mostly to the prototyping aspect of construction, and in part to an inefficient use of materials. In most instances, robotics research has only sought to duplicate human labor without questioning the process in the first place. As a result, research has lead to islands of automation with ever larger machines intended to assemble ever larger structures. Related but not negligeable is the inefficient use of material.

As a result, research has lead to islands of automation with ever larger machines intended to assemble ever larger structures. Related but not negligeable is the inefficient use of material. The insane amount of solid waste generated by the construction industry may contribute as much as 23% of the national stream [Apotheker, 1990].In view of the similarities between construction and solid freeform fabrication, our research has taken the opposite route. Instead of giant machines assembling giant parts, a large structure could conceivably be built by an army of ants, one grain of sand at a time.3 EXPERIMENTAL PROCESS As pointed out in Section 1, selective aggregation of bulk solid material appears to be the most likely technique capable of achieving parts of meter size and over.

Selective aggregation can be effected in the following manners: Selective deposition of an activated binder into the bulk mate rial, as in the 3-D Printing process [Sachs et aI., 1990]; selective activation of a passive binder present in the bulk material -including self-cementing materials- as in the Selective Laser Sin tering Process [Deckard et aI., 1987]; or activation of a binder selectively deposited into the bulk material. The latter was chosen as it is also the most flexible in permitting fabrication from a very wide range of bulk materials in various forms (powder, pebbles, or blocks for example.)As is now customary in Solid Freeform Fabrication, the basic principle of operation is to incrementally build-up layers of the structure. The geometry of each layer is obtained directlyFIGURE 1.FUNCTIONAL DIAGRAM OF LAYERED CONSTRUCTION PROCESS.CAD SYSTEM41from slicing a geometric model and is translated into scanning motion of the end-effector.

This process is illustrated by the diagram in Figure 1 and was adopted for this project.The core of the process involves selective deposition of a powdered binder material onto a bed of powdered matrix material. From a material handling perspective, this technique is akin to the age old tradition of sand painting practiced for centuries by Navajo Indians and Tibetan Monks. The resulting mix is then activated by a catalyst (in this case water). The process is illustrated by Figure 2.For lack of funds, the approach of automated deposition suggested in Figure 2 was not imple mented. Instead, masking processes were combined with uniform deposition of matrix and binder to achieve the desired result. The masks were obtained by slicing the geometry and plotting the outer and inner layer boundaries on two different masks. The two masks were used in combina tion to sift the binder material and then removed leaving only the exposed cross-section impreg nated with binder.

The two masks were used in combina tion to sift the binder material and then removed leaving only the exposed cross-section impreg nated with binder. The process steps are illustrated in Figure 3.The material used for this experiment were sand and Portland Cement. A series of sample hol low concrete structures 3"x3"x6" was fabricated in layers. The geometry of the sample is shown in Figure 4. Note that this structure could not be produced by regular concrete casting techniques. Various tests were conducted with and without compressing the granular sand and cement layer. Various sand to cement ratios were experimented with. Wetting of the layer was initially per formed by spray mist of liquid water onto each layer as it was completed and cured. It was found that this approach led to brittle samples. When observed under the microscope, it was seen that cement had formed large aggregates and had not bonded with sand.

It was found that this approach led to brittle samples. When observed under the microscope, it was seen that cement had formed large aggregates and had not bonded with sand. Regardless of the sand to cement and water ratios, the concrete structure had cracks running along the layer structures. This phenomenon is attributed to the effect of surface tension in the deposited droplets of liquid water.Excellent bonding was obtained however when steam was used in place of water. After all lay ers were completed, the sample was placed into a pressurized steam chamber. The pressure and temperature in the chamber were respectively 3 atmospheres and 300°C and caused the concrete to cure rapidly. When observed under a microscope, the layers were so well bonded that it wasFIGURE 2.FLOW DIAGRAM OF THE SELECTIVE DEPOSITION PROCESSl r - - -....

When observed under a microscope, the layers were so well bonded that it wasFIGURE 2.FLOW DIAGRAM OF THE SELECTIVE DEPOSITION PROCESSl r - - -.... CD ~® ::z-:=-a0 ----,Blanket deposition of matrix materialSelective deposition of reactive agentActivation of the reactive agent42impossible to tell them apart. The concrete obtained by this method is dense and has no apparent porosity or cracks. Vibration tests were also conducted using a bench designed for testing elec tronic equipment at Teledyne Corporation. These vibration tests, which destroyed all the water sprayed samples, failed to break the steamed samples. Later, tests were conducted applying vapor at atmospheric pressure in between each layers and gave similar results. A picture of the part is shown in Figure 5.The choice of materials used in this experiment was justified by Indeed cost considerations alone. . . - - - - - - - - - - - - - - - - - - - - - - . . ..

A picture of the part is shown in Figure 5.The choice of materials used in this experiment was justified by Indeed cost considerations alone. . . - - - - - - - - - - - - - - - - - - - - - - . . . . , the proposed process is essentially about selective deposition of pow ders that either react together when exposed to a catalyst (local ized heat, electric current) or a chemical (water in this case.)FIGURE 3.FLOW DIAGRAM OF THE SELECTIVE DEPOSITION PROCESSBlanket deposition of matrix material~ ®Masking ~ oBlanket deposition of reac- ~ oMask and excess reactive y ®Activation of reactivetive bulk material.material removal.4 MATERIAL CHARACTERIZA-TIONThis material was not available in time for publication of this paper and will be presented at the sym posium.5 TECHNOLOGY ASSESSMENT Many construction processes qual ify as additive manufacturing. This is what may have caused the demise of robotics since the part and the robot's environment are one and the same.

This is what may have caused the demise of robotics since the part and the robot's environment are one and the same. Manufacturing automation usually assumes a con stant environment -e.g. plant or workcell- and changing parts but it does lack sensor and infor mation processing capable of deal ing reliably with a changing environment. This is where solid freeform fabrication has a definite advantage over robotics. Indeed, in SFF, the part is incrementally built in an organized manner. We claim that this attribute makes SFF ideally suited for off-site as well as on-site construction.material.43FIGURE 4. SPECIFICATIONSSAMPLE3 in.TEST5.1 Rapid Prototyping of Large Functional StructuresPA'ihd Off-Site Construction~Solid Freeform Fabrication has already spun off number of successful applications in rapid prototyping, rapid tooling and mold making.

To be useful for large parts and off-site construction of large functional structures, however the process must reach volumetric flowrates that make SFF economically viable.c::: ~ Attempts are underway at scaling up existing pro- S cesses to produce large scale prototypes with character ~ istic sizes of a meter and above. However, current SFF technology hits a technological barrier when it comes to ;. ~ increasing the volumetric flow rate of build material.en enFIGURE 5. BLOCK-HOUSE LAYERED STEAMED CEMENT-SILICAHO PRODUCED MANUFACTURINGSAMPLEAs per the classification presented in Section 2, the maximum volumetric flow rate is obtained when the build material is already in bulk solid form, leaving selective aggregation as the only viable alternative for Bylarge parts. Another aspect of the same problem IS the dt9fvery of binder Binder material should deposited at a rate that satisfies both the build flow rate and the geometnc accu- racy requirements.

Another aspect of the same problem IS the dt9fvery of binder Binder material should deposited at a rate that satisfies both the build flow rate and the geometnc accu- racy requirements. Our process achieves this result with deposition of the binder in powder form, leaving activation as a blanket process.SCALE..5.2 On-Site Construction AutomationWe argues that layered fabrication is a viable approach to on-site automated construction. In a first phase, our technol ogy is expected to be applicable to the automated construc tion of simple on-site structures such as catch basin, foundations, and decks.As the technology develops, it may also become a viable commercial alternative to lumber construction. Because this technology offers on-site construction automation, it would reduce the dependence upon labor, the risk of injuries, and weather stoppage.

Because this technology offers on-site construction automation, it would reduce the dependence upon labor, the risk of injuries, and weather stoppage. A rough estimate is that this technology may cut construction cost by about 30%.An advantage of the proposed approach to construction automation is its environmental bene fits. A large body of work has demonstrated the effectiveness of using waste material, such as fly ash as a construction material [Goumans et aI., 1994]. Our general purpose material handling approach makes using such materials a viable alternative in the process. A further benefit of SFF is its thrifty use of material. Indeed, all unused material can be readily recycled, contributing to a significant waste reduction in construction.44.6 CONCLUSION Proof of concept for a new process aimed at solid freeform fabrication of large (> 1meter) struc tures was performed.

In this paper, we presented a characterization of additive manufacturing pro cesses and argued that selective aggregation was the only technique prone to fabrication of large scale structures. To prove the point, a novel system of layered fabrication was studied. It involved depositing a layer of reactive material (Portland cement) over a layer of matrix material (silica). Two methods were proposed to perform the task of binding the cross-sectional material: (1) the binder is selectively deposited and blanket activated, or (2) the binder is blanket deposited and selectively activated. Parts were built out of sand (matrix) and Portland cement (reactive material) activated by water vapor. The resulting structure is denser than, and exhibits mechanical proper ties similar to cast concrete.7 BIBLIOGRAPHY [Apotheker,1990] Apotheker, S., "Construction and Demolition Debris-The Invisible Waste Stream," Resour. Recycling, 9(12), pp.

The matrix material of choice was nylon 6. Samples were printed exploring a range of reinforcement volume content. In summary, this study presents an assessment of the applicability of through-thickness testing to AM CFRP specimens and provides a performance comparison between AM composite through-thickness properties and the properties of equivalent CM CFRP specimens.Keywords carbon fibre, through thickness, composite, mechanical behaviour, additive manufacture1. Introduction1.1. Composite materialsComposite materials are non-homogenous materials made from two or more constituent materials that differ in their mechanical and chemical properties and provide desirable tailored characteristics when combined. Fibre-reinforced polymers (FRPs) are composite materials made of a polymer matrix, which houses, supports and transfers the bearing load onto a fibre reinforcement that acts as the load-bearing structure in the composite.

The use and manufacture of FRPs has been widespread since the 1930s, with the mass production of glass fibres discovered in 1933 [1] and they are today commonly used in industries that include aerospace, automotive, construction and defence [2].1.2. Carbon fibre reinforced polymersCarbon fibre reinforced polymers (CFRPs) use carbon-based (usually graphite) fibres of 5 µm to 10 µm in diameter, twisted into yarns that are woven or embedded into a polymer matrix [3]. These fibres provide good tensile properties with respect to their mass, making them useful for a large number of applications, including aerospace, marine and automotive parts, as well as end-user products, such as golf clubs and fishing rods.1.3. Through-thickness testing of CFRPsThe nature of CFRPs and other matrix-based composites, where fibre reinforcement is applied in a plane-wise fashion, that is, in a layer-by-layer format, means that CFRP structures exhibit anisotropic properties [4].

This is primarily due to the relatively low structural strength of the matrix (polymer) material of the composite, as well as, in some cases, a relatively weak616fibre-matrix interface between layers, which may constitute a bottle-neck in the transverse (Z- direction) structural strength of a CFRP structure [2].Through-thickness testing (also known as Z-direction tensile testing) is a mechanical property assessment method that applies a tensile force to a test specimen whose axis is perpendicular to the reinforcement plane of a composite material (Figure 1) [5]. The main purpose of through-thickness testing is to assess the resistance to lamellar tearing [6], a form of (initially microscopic) cracks that form parallel to the plate surface of a material and induce weaknesses in the material’s through-thickness (Z) direction.

The process of through-thickness testing is well-understood in the polymer composite world, and specification standards have been developed for testing matrix composite materials in the through-thickness direction [5]. Several studies have been conducted on through-thickness performance testing of matrix composites [7][8][9] and some introduced additional Z-direction reinforcements before testing [10]. However, these studies focus on the conventionally processed composites, but no literature is available on through-thickness testing of additively-manufactured (AM) composites.Figure 1: Through-thickness tensile test specimenThis study will aim to shed light on the applicability of through-thickness testing on AM parts as well as to evaluate their performance. This is of particular importance since AM is the latest process for manufacturing complex, non-homogeneous, micro-structured parts such as CFRPs.

This is of particular importance since AM is the latest process for manufacturing complex, non-homogeneous, micro-structured parts such as CFRPs. Current composite manufacturing processes, including thermoplastic compression moulding, filament winding, resin infusion, etc., offer various features, advantages and disadvantages, which provide variety in approaches and the potential to optimise production according to a balance of needs (e.g. size, quality, production rate, tooling costs). AM is the latest composite manufacturing process to be added to the list, which comes with its own design opportunities and challenges. Therefore, it is important to establish AM’s617properties by applying relevant test methods to specimens produced by it, adding to the body of knowledge on the different properties that this new composite manufacturing process can bring forth.1.4.

CFRPs in AMWhile conventional manufacturing techniques of CFRPs have been well-studied, the long cycle times and high production costs associated with their production have given AM a distinct potential advantage over other manufacturing processes, because AM carries the potential for both reducing time and cost of production as well as increasing design complexity of the end products [11].The two main AM techniques that have been studied for the production of CFRPs are powder bed fusion, e.g. selective laser sintering (SLS) [12][13], and material extrusion, e.g. fused deposition modelling (FDM) [14][15] – also known as fused filament fabrication – with the majority of research into polymer composites falling into the latter method [16][17][18].1.5.

fused deposition modelling (FDM) [14][15] – also known as fused filament fabrication – with the majority of research into polymer composites falling into the latter method [16][17][18].1.5. Material extrusion of CFRPsMaterial extrusion (ME) (Figure 2) is an AM process that makes use of polymer filament feedstocks by extruding the material in its melted form, through a computer-controlled liquefier head, into a build platform in a layer by layer fashion. Advanced ME machines can support multiple extrusion nozzles with multiple materials. For example, in this case, a polymer matrix material and a separate fibre reinforcement material. Several studies have tested ME composites compared to composites manufactured using conventional methods and have shown promising results in tensile strength [15][16].

Several studies have tested ME composites compared to composites manufactured using conventional methods and have shown promising results in tensile strength [15][16]. While several studies into the tensile properties of AM composite parts have been conducted [14][15][16], these studies focused on the longitudinal (in-plane) tensile properties, rather than the transverse (through-thickness) tensile strength.2. Materials and methods2.1. Manufacturing equipment and parametersThe experiments utilise an in-house Markforged Mark Two material extrusion machine for manufacturing the test specimens. The machine uses fused filament fabrication (FFF) and composite filament fabrication (CFF) technologies to manufacture functional polymer and composite polymer parts [19].

The machine uses fused filament fabrication (FFF) and composite filament fabrication (CFF) technologies to manufacture functional polymer and composite polymer parts [19]. The machine is capable of printing four types of reinforcement material: carbon fibre, fibreglass, Kevlar and high strength, high temperature (HSHT) fibreglass, in addition to two choices of matrix material: nylon and the Markforged Onyx [20].Test specimens were manufactured from nylon 6 as a matrix material and carbon fibre as reinforcement. All materials used were Markforged-certified materials and were obtained from CREAT3D Limited, a certified local distributor.The print bed was levelled and print nozzle heights adjusted prior to the first print. The print bed was coated with a layer of Elmer’s washable school glue in order to protect the bed, facilitate part removal and prevent the part from moving while the print is in progress.

The print bed was coated with a layer of Elmer’s washable school glue in order to protect the bed, facilitate part removal and prevent the part from moving while the print is in progress. A 6 mm- wide nylon brim was printed under each specimen in order to increase the contact area with the print bed and prevent thermal curling. The nylon and carbon fibre nozzle temperatures were set to 275 °C and 245 °C, respectively.618Figure 2: Schematic of the material extrusion process.Nylon support structures were used during the printing process, which were later removed through manual prying with pliers. A purge line was printed prior to each print in order to purge the print head and extruder from any nylon that may have accumulated moisture from the atmosphere during idle times. The print head was also purged of nylon residue that may have accumulated on the nylon nozzle while carbon fibre was routed through the part.

The print head was also purged of nylon residue that may have accumulated on the nylon nozzle while carbon fibre was routed through the part. All specimens were printed with 100% fill density with a rectangular fill pattern to ensure the maximum fill density as per manufacturer recommendations. Layer height was fixed by the manufacturer at 0.125 mm. The specimens were built five at a time in the spacing shown in Figure 3.2.2. Specimen designAll test specimens were manufactured to comply with ASTM D7291/D7291M-15 standards on through-thickness tensile testing [5]. The specimens were manufactured in accordance with recommendations from the National Physical Laboratory (NPL) [21], as shown in Figure 4. The test piece is a standard Royal Armament Research and Development Establishment (RARDE) specimen with 25 mm square ends and a 40 mm nominal height. The reduced gauge section is 12 mm long and has a rectangular cross-section of 10 mm × 16 mm.

The reduced gauge section is 12 mm long and has a rectangular cross-section of 10 mm × 16 mm. The reinforcement was applied in the x-y plane (perpendicular to the height of the specimen) and was applied uniformly along the entire height of the specimen.The reinforcement in each layer was laid out in a concentric fashion, where carbon- fibre rings ran continuously from the perimeter to the centre (Figure 5). This amounted to twelve rings in base layers and three rings in middle layers.619Figure 3: Specimen spacing in each printFigure 4: Schematic of through-thickness tensile CFRP specimen in accordance with ASTM D7291/D7291M-15 standards and NPL recommendations [21]Figure 5: Per-layer layout of carbon fibre rings (blue) within nylon matrix and supports (white) at the base (right) and in the gauge section (left)6202.3.

Specimen sets and experimental designFour sample sets were produced for through-thickness tensile testing, each set consisting of five identical specimens to provide statistical significance and room for printing errors. The four sets were as follows:1. Control samples: 100% nylon samples, no composite material deposited. These were used to benchmark the rest of the samples and provide a comparative baseline.2. 25% fibre volume fraction (fvf) of carbon fibre (one layer of fibre in every three layers of nylon), evenly distributed along the through height of the specimen, i.e. homogeneously reinforced.3. Half-density samples: approx. 37% fvf of carbon fibre to further test for correlation between fibre content and tensile strength.4. Full-density samples: i.e. composite reinforcement present in print every layer, which amounts to approximately 78% fvf of carbon fibre for maximum possible fibre density.2.4.

Full-density samples: i.e. composite reinforcement present in print every layer, which amounts to approximately 78% fvf of carbon fibre for maximum possible fibre density.2.4. Sample preparationSupport structures were removed from the printed samples using pliers. The test samples were then measured to check machine dimensional accuracy and prepared for tensile testing in accordance with ASTM D7291/D7291M-15 standards. As part of the preparation, a number of end tabs were machined from 6082 aluminium alloy bars using a XYZ Proturn SLX 425 lathe. End tabs and specimens were later grit-blasted using 150 µm to 300 µm glass beads in order to provide a suitable surface for adhesive bonding.A number of suitable adhesives were tested to determine the best choice for bonding the samples to end tabs. The Loctite 410 instant adhesive was chosen. Specimens were bonded to the aluminium alloy end tabs using metal jigs to align the parts.

The Loctite 410 instant adhesive was chosen. Specimens were bonded to the aluminium alloy end tabs using metal jigs to align the parts. Specimens were then left to cure for a minimum of twenty-four hours prior to testing.Following curing, specimens were sprayed with a black and white texture in order to allow for the video strain gauge to record specimen strain under tensile loading while testing (Figure 6). A total of four specimens were bonded and prepared for the first round of testing (one of each set), followed by a second round for the remaining sixteen specimens.2.5. Tensile testing and analysisAn Instron 5965 Mechanical Test Machine was used to determine the tensile properties of the specimens. This was then to be used to perform an evaluation on through-thickness properties and ultimately the applicability of through-thickness testing to AM parts.

This was then to be used to perform an evaluation on through-thickness properties and ultimately the applicability of through-thickness testing to AM parts. The strain under tensile loading was measured using an iMetrum video strain gauge that tracks the displacement of a set of texture dots painted on the specimens.621Figure 6: Through-thickness tensile specimen with fitted end-tabs and texture spray paint ready for testing.The test was conducted at 18 °C and 50% relative humidity. Due to time and resource restrictions, a test speed of 1mm/min was used as per NPL testing protocols [22] instead of the ASTM D7291/D7291M-15 0.1 mm/min.3. Results and discussion 3.1. ManufacturabilityA total of five failed and twenty successful specimens were produced during the printing process in addition to a single-sample test print. The single pure nylon sample was printed in order to ensure all parameters are set correctly and that no complications during the print process occurred.

The single pure nylon sample was printed in order to ensure all parameters are set correctly and that no complications during the print process occurred. Following the successful test print, five Set 1 (0% carbon fibre) specimens were added to the print job. However, significant distortion was present on all five samples (Figure 7). The distortion observed on all five samples is believed to be the result of thermal curling induced by the rapid cooling of nylon layers as they were laid onto the unheated substrate. Adding a brim to the parts seemed to reduce thermal distortion but failed to eliminate it. It was later observed that the addition of carbon fibre to the specimens significantly increased part stiffness and dimensional integrity. As a result, the pure nylon parts were re-printed with fibre-reinforced bases and a pure nylon gauge (Figure 8).

As a result, the pure nylon parts were re-printed with fibre-reinforced bases and a pure nylon gauge (Figure 8). Additionally, prior to each print, the print substrate was put under running hot tap water until hot in order to reduce the thermal gap between the printer nozzle and the substrate.622Figure 7: Thermal distortion in original Set 1 specimens.Figure 8: Set 1 (0% carbon fibre) specimen with reinforced bases.Thermal distortion was only observed during the five-sample print - but not in the initial single-sample print - despite both being printed with pure nylon. The reason behind this is believed to be the five-fold delay in laying each consecutive layer onto the previous ones as the machine continued to build the next layers on the remaining four specimens, which allowed more time for the plastic to cool down and deform. The option to programme the machine to print the samples in succession, rather than simultaneously, was considered.

The option to programme the machine to print the samples in succession, rather than simultaneously, was considered. However, as of the time of writing, the printer firmware was incapable of incorporating such measures.623The production of Set 2 (25% fvf) and Set 3 (37% fvf) samples was successful. Set 4 (78% CF) prints showed carbon fibre splinters protruding through the matrix throughout the specimen’s height (Figure 9). Repeat prints of Set 4 specimens conducted by the manufacturer confirmed the observation. The reason for this observation is thought to be excessive fibre content, where the density of fibre in the part is too great for the printer to print successfully. This may have caused some of the printed fibre to either tangle with previously-laid fibres or to dislodge the fibre, causing splinters and possible breaks in the printed fibre. In applications where maximum density is required, it is recommended that the parts be ‘striped’.

In applications where maximum density is required, it is recommended that the parts be ‘striped’. That is, to print several layers of fibre in succession (usually ten) followed by a few (usually two) pure nylon layers. This is thought to mitigate the effect of fibre splintering. This has not yet been done, but it is planned for future investigation.Figure 9: Set 4 specimen with CF splintersDimensional measurements were recorded for the gauge sections of each specimen using a calibrated Clarke® Electronic Digital Calliper. A total of ten repeats were taken of each measurement and the tolerances were calculated with 95% confidence. As shown in Table 1, fibre-reinforced specimens consistently showed closer values to the target dimensions of 10 mm and 16 mm of the gauge section than the non-reinforced samples.

As shown in Table 1, fibre-reinforced specimens consistently showed closer values to the target dimensions of 10 mm and 16 mm of the gauge section than the non-reinforced samples. This is primarily due to the fact that carbon fibre adds stiffness to the printed parts, thus greatly reducing the effects of the various factors that may influence the dimensional integrity, such as thermal and gravitational deformations. Conversely, fibre-reinforced specimens generally exhibit larger tolerances than non-reinforced samples.

Test resultsEight out of twenty specimens failed due to premature adhesive failure. The average stress value reached at the time the adhesive failed resulting in the tensile test being terminated was 7 MPa ± 1.5 MPa. Several factors could have played part in the failure of the adhesive. Firstly, since samples were left to cure for a total of twenty-four hours prior to testing, it is plausible that the adhesive had not reached its ultimate strength and that further time should have been allowed for sufficient curing. Secondly, it is possible that the specimens may have required a coarser surface texture in order to provide a larger surface area for bonding. A good practice guide by the Composites UK Trade Association on the adhesive bonding of composites shows that medium grit (120 µm to 200 µm) is ideal for composite surface pre-treatment.

A good practice guide by the Composites UK Trade Association on the adhesive bonding of composites shows that medium grit (120 µm to 200 µm) is ideal for composite surface pre-treatment. The guide further recommends that cryo-blast and soda-blast be used in treating composites, as opposed to shot-blasting, which is more suitable for metals [23]. Indeed, upon touch, little difference in surface roughness was observed following the grit-blasting of the specimens than the original 3D-printed surface roughness. A more suitable grit will be considered for future trials. Finally, nylon is notoriously difficult to bond due to the presence of slip additives within the material, which can migrate to the surface of the material and interfere with proper bonding. As a result, it may have been difficult for the adhesive to properly bond to the nylon surface, thus causing premature delamination in the glue and nylon interface.Successful specimens were as follows:1.

Set 1 (0% fvf): Five out of five specimens successful. 2. Set 2 (25% fvf): Two out of five specimens successful. 3. Set 3 (37% fvf): One out of five specimens successful. 4. Set 4 (78% fvf): Five out of five specimens successful.6253.3. Fracture surfaces Set 1 specimens All five Set 1 specimens fractured at the interface between the fibre-reinforced base and the start of the gauged section, as opposed to the expected typical fracture within the gauge section itself. This suggests apparent weakness in the adhesion between carbon-fibre- reinforced layers and unreinforced layers. The cause of this systemic weakness is unknown, but it is thought that the sudden drop in fibre volume fraction may be inducing certain defects in the interface within the printed part. The fracture interface itself (Figure 10) exhibited ductile behaviour in all five specimens.

The fracture interface itself (Figure 10) exhibited ductile behaviour in all five specimens. This is due to the ductile nature of the nylon matrix.Figure 10: The ductile fracture interface of Set 1 specimensSets 2 and 3 specimens The fracture surfaces of successful specimens from Sets 2 and 3 were similar. All successful specimens belonging to the two sets failed within the gauge section. The fracture interface was horizontal and suggested a final matrix failure (Figure 11). It was clear that the fibre had some function in resisting tension, as fibres were observed to stretch and break during testing.626Figure 11: The fracture interface of Sets 2 and 3. The top photo shows carbon fibre being stretched during testing, suggesting the fibre assisted in resisting tensionSet 4 specimens Similarly to Sets 2 and 3, the failure occurred within the gauged section for all successful specimens.

However, the role of carbon fibre providing tensile strength was much more apparent in Set 4 specimens as can be seen from the stretched and torn fibres embedded in the fracture surfaces (Figure 12).Figure 12: The fracture interface of Set 4 specimens. Torn fibres suggest prominent role of CF in providing tensile strength6273.4.

These graphs only illustrate specimens that have completed the test successfully.As shown in Figure 13, the tensile behaviour of Set 1 specimens matched the ductile fracture patterns observed. While most samples had a fairly consistent elastic behaviour, the variation in UTS was high. This phenomenon can be explained by the possible presence of random print defects, which can unpredictably alter the tensile strength of the specimen. These defects are likely to be due to the sudden change in fvf. More research is required to study the effects of changes in fvf on part integrity. All specimens exhibited typical ductile behaviour, including the presence of high and low yield points. The ductile behaviour is consistent with the expected behaviour of nylon 6.Figure 14 illustrates the tensile behaviour of the specimens from Sets 2 and 3. They exhibit consistent elastic behaviour and recorded the highest UTS values.

They exhibit consistent elastic behaviour and recorded the highest UTS values. This suggests the existence of an optimum fvf at which the combined strength of the nylon matrix and carbon fibre yields the highest tensile strength. Further research is required to determine this optimum point. The specimens were significantly less ductile than Set 1 specimens in their post-yield behaviour. This is expected as the carbon fibre provides additional stiffness to the specimens. However, further tests are required to provide better certainty of the results.628The tensile behaviour of Set 4 specimens is shown in Figure 15. As with other reinforced specimens, the elastic behaviour of all specimens tested was consistent, as were the UTS values. This is expected as the carbon fibre provides a more consistent build. However, the UTS values of Set 4 specimens were significantly lower than those of Sets 2 and 3.

This is expected as the carbon fibre provides a more consistent build. However, the UTS values of Set 4 specimens were significantly lower than those of Sets 2 and 3. This suggests that the higher fvf came at the expense of matrix material, whose primary function is to provide interlayer adhesion, thus reducing overall strength.Figure 13: Tensile stress behaviour of Set 1 specimensFigure 14: Tensile stress behaviour of Sets 2 and 3 specimens629Figure 15: Tensile stress behaviour of Set 4 specimens3.5. Future WorkFuture work is planned to study the effects of print defects, fvf variations and conductmore thorough through-thickness tensile testing.More specimensMore specimens will be manufactured in order to study the effect of the sudden change in fvf. This may be accompanied with SEM inspection of fracture interfaces to detect possible print defects in the interface between the last reinforcement layer and the first pure nylon layer.

This may be accompanied with SEM inspection of fracture interfaces to detect possible print defects in the interface between the last reinforcement layer and the first pure nylon layer. Different measures will be explored in order to prevent thermal deformation of unreinforced samples, including the introduction of a thick raft to be filed down in post-process.Additional sets exploring a range of reinforcement ratios will be printed in order to determine the optimum fvf that provides the highest UTS values. All sets would have at least ten specimens in order to provide additional statistical significance as well as allow for failed parts.Optimisation of the testAs demonstrated through the trials, despite following proper protocols, specimen bonding results were erratic, where glue failure occurred on a range of loads which seemed to follow no particular pattern. This is largely due to the difficult nature of nylon as a low surface energy material.

This is largely due to the difficult nature of nylon as a low surface energy material. As such, options will be explored in adjusting and altering the test method in order to account for inherent difficulties in both the material and the process, including exploring the options to bond specimens using a lap joint arrangement.AcknowledgmentsThe authors would like to thank Dr Graham Sims, Dr Richard Shaw and Dr PeterWoolliams from the National Physical Laboratory for supporting this work.630FundingThis work was funded by EPSRC [Grant EP/L01534X/1], Exova and the NationalPhysical Laboratory.4. References[1] G. Slayter, Method and apparatus for making glass wool. US2133235 A, 1938.[2] M. Piggott, Load bearing fibre composites, Kluwer Academic Publishers, Dordrecht, 2005.[3] Zoltek, What is Carbon Fiber?. http://zoltek.com/carbonfiber/, 2017 (accessed 08.06.17).[4] F.

PLICATON OF DESIGN OF EXPERIMENTS (DOE) ON THE PROCESSING OF RAPID PROTOTYPED SAMPLESJ.Weinmann, H. Ip, D. Prigozhin, E. Escobar, M. Mendelson, and R. NooraniNational Science Foundation Research Experience for the Undergraduates Program Loyola Marymount University Los Angeles, CAAbstract The purpose of this experiment was to improve the Fused Deposition Modeling Process by examining the tensile strength of samples fabricated in a Stratasys FDM 1650 Machine utilizing the methods of Design of Experiments. A two-level, four-factor, full factorial experiment was conducted. The selected factors were temperature, air gap, slice thickness, and raster orientation. A regression equation determined the level each factor should be set in order to optimize the FDM machine settings.

A regression equation determined the level each factor should be set in order to optimize the FDM machine settings. It was found that single factors - small air gap, small layer thickness and low raster orientation, as well as the interaction between high temperature and small layer thickness yielded the greatest effect the response.1. IntroductionFused deposition modeling (FDM) is a process that is used for fabricating solid prototypes from a computer-aided design (CAD) data file [1]. The process fabricates 3-D parts from a build-up of 2-D layers. In this process, an Acrylonitrile-Butadiene-Styrene (ABS) thermoplastic polymer is extruded through a heated nozzle to deposit the layers. In a previous paper, we showed that the orientation of the layers created anisotropic tensile properties [2].

In a previous paper, we showed that the orientation of the layers created anisotropic tensile properties [2]. Other investigators have also experienced similar results [3].Previous investigators have used design of experiments (DOE) as a method to maximize strengths of the FDM-processed specimens of silicon nitride [4] and ABS polymer [3]. Many of the factors used for influencing the strength were entirely different in these studies. These investigators have not physically interpreted their selected factor levels in terms of the material properties and microstructure.The purpose of this paper is to use quality engineering tools to design, analyze andphysically interpret our selection of the FDM processing factors and their levels.2. Experimental MethodsThe 12-step design process was used for our experimental [5].

First, the problem of concern was the low strength of FDM-processed ABS test specimens; second, our objective was to maximize the tensile strength; third, the yield and ultimate strengths were selected as the quality characteristics, i.e., the response; fourth, the factors were determined by team brainstorming and were recorded on a cause-effect diagram [6]; fifth, four factors and no-noise factors were selected for our study; and sixth, a two-level experiment was selected.340341Y yield (MPa) 268˚C -0.0254 mm 0.254 mm 0˚/45˚ 15.70 268˚C -0.0254 mm 0.254 mm 90˚/45˚ 12.62 268˚C -0.0254 mm 0.356 mm 0˚/45˚ 11.99 268˚C -0.0254 mm 0.356 mm 90˚/45˚ 9.06 0.254 mm 0˚/45˚ 12.97 268˚C 0.254 mm 90˚/45˚ 12.98 268˚C 0.356 mm 0˚/45˚ 11.91 268˚C 268˚C 0.356 mm 90˚/45˚ 7.67 277˚C -0.0254 mm 0.254 mm 0˚/45˚ 17.56 277˚C -0.0254 mm 0.254 mm 90˚/45˚ 18.70 277˚C -0.0254 mm 0.356 mm 0˚/45˚ 10.31 277˚C -0.0254 mm 0.356 mm 90˚/45˚ 9.17 0.254 mm 0˚/45˚ 13.70 277˚C 0.254 mm 90˚/45˚ 12.53 277˚C 0.356 mm 0˚/45˚ 10.60 277˚C 277˚C 0.356 mm 90˚/45˚ 6.67 Table 2.

L16 (24) Experimental Design ReportTrial ABCD1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 160 mm 0 mm 0 mm 0 mm0 mm 0 mm 0 mm 0 mmY ultimate (MPa) 16.31 13.64 12.86 10.36 13.38 13.81 12.83 8.77 17.77 19.00 11.39 10.35 13.90 13.64 11.36 8.834. Analysis of Data Based on the results in Table 2, the column effects for factors A, B, C, D and their interactions were determined for both yield and ultimate strengths. The analysis of the effects on the response was plotted in Figs. 1 and 2. The ~95% (±2σ) confidence interval for the standard error of the effects [8] of each factor was calculated to be ±0.73 MPa. When the response lies outside this interval, the effect is significant; and the converse is true when the effects are inside the interval. In Fig.

When the response lies outside this interval, the effect is significant; and the converse is true when the effects are inside the interval. In Fig. 1, factors B (air gap), C (layer thickness), D (raster orientation), and interaction AC lay outside the 95% confidence interval 12.13 ± 0.73 MPa, which indicates only these factors and interaction have significant effect on yield strength. Fig. 2 shows the detailed analysis of the AC interaction. The conclusion from these figures is that to maximize the yield strength, the low- level of factors B, C and D, and the interaction of high-level A and low-level C should be selected.34215.0014.50C -114.0013.50s e u a V e s n o p s e R e g a r e v Al13.0012.5012.0011.5011.00A +1A -1B -1B +1D -1D +1AB -1AB +1AC -1AC +1AD +1AD -1BC +1BC -1BD -1BD +1CD -1CD +1ABC +1ABD -1ABC -1 ABD +1BCD -1 ACD -1 ACD +1 BCD +110.5010.009.50C +19.00Fig. 1.

3. Ultimate Strength Response (MPa) vs. Effects of Factors/LevelsIn order to verify the above results, a regression equation was used to predict differentfactor levels and two-factor interactions that maximized the strength response, (cid:468) [9]:(cid:468) = βo + Σi βi xi + ΣiΣj βij xI*x j = βo + βB xB + βC xC + βD xD + βAC xA+1*xc-1(1)Here xi represents factors B, C and D which take on ±1 values for the low/high levels that will maximize tensile strength. Also, xi*xj represents the interaction of A+1*C-1 that takes on ±1 values. The values of the β coefficients are shown in Table 3. Coefficient βo represents the intercept (or mean value of Y). Coefficients βi are the slope of multiple regression for factors B, C and D. βAC represents the coefficient that is associated with the A*C interaction. Here the coefficients are one-half the effects of each factor and interaction in Figs. 1 and 3.

βAC represents the coefficient that is associated with the A*C interaction. Here the coefficients are one-half the effects of each factor and interaction in Figs. 1 and 3. The values of the coefficients are shown in Table 3.Tensile Strength (MPa) Yield Ultimateβ0 12.13 13.01βΒ -1.01 -0.95Coefficients βC -2.46 -2.17βD -0.96 --βAC -0.76 --Table 3. Calculated Coefficients for B, C, D Factors and Interaction AC in Eq. (1).Eq. (1) verifies that for maximum yield strength, factor levels should be B-1, C-1, D-1, andA+1C-1. For maximum ultimate strength, the factor levels should also be B-1 and C-1.When the interaction effects are moderate (see Fig. 2), errors can result in selecting the factor levels if the interactions are not taken into account. For example, in previous work [4], all of the columns in their 3-level experimental design were filled with main factors.

For example, in previous work [4], all of the columns in their 3-level experimental design were filled with main factors. This created a lower resolution experimental design, where the interactions and their aliases could not be analyzed. In this case, Eq. (1) would become an additive response of main factors, where ΣiΣj βij xi*xj = 0. When our interactions were analyzed, A+1C-1 was preferred over A-1C-1.3445. Physical InterpretationThe selected process factor levels must physically make sense from the point of view of the structural properties and microstructure of ABS polymer. The yield strength was correlated with stretching the polymer chains and viscoelastic flow of ABS, and the ultimate strength was correlated with fracture initiation through the ABS structure [7].

The yield strength was correlated with stretching the polymer chains and viscoelastic flow of ABS, and the ultimate strength was correlated with fracture initiation through the ABS structure [7]. Once a critical crack length was initiated, it propagated either at 45° or 90° to the tensile axis depending upon the raster orientation.The FDM machine deposited the raster patterns at 0°/45° and 90°/45° oriented composite structures. The weak interfacial bonding between the oriented fibers can be caused by (1) weak interlaminar shear properties of the fibers, (2) the volume change during the ABS transformation from liquid to solid, or (3) formation of pores during FDM processing [2].Factor B – Air Gap: When the air gap was set at a negative value, the adjacent fibers overlapped each other. This increased the bonding between the fibers and created a tighter structure, as the porosity between the fibers was reduced.

This increased the bonding between the fibers and created a tighter structure, as the porosity between the fibers was reduced. The lower porosity composite is expected to translate into a higher tensile strength. Hence, it is reasonable that a lower air gap setting (B-1) would maximize the composite strength.Factor C – Layer Thickness: When the layer thickness was reduced, the fiber diameter was reduced, and the fiber shape became more oval as shown in Fig. 5. Also, the 0o fibers (light phase) overlapped each other to a greater extent when the layer thickness was low. The lower layer thickness reduced the porosity and increased the volume fraction of fibers (to a smaller extent), which strengthened the overall composite structure.

The lower layer thickness reduced the porosity and increased the volume fraction of fibers (to a smaller extent), which strengthened the overall composite structure. This adequately explains why a low layer thickness (C-1) is selected for increasing the tensile strength.Factor D – Raster Orientation: When the two raster orientations were compared, the difference between them was in the 0o and 90o oriented fibers since the 45o layers acted similarly in both cases. Fibers oriented parallel to the tensile axis would exhibit maximum strength, while those oriented perpendicular to it would have their weakest strength. This is shown in Fig. 4 where the 90o/45o structure fibers perpendicular to the tensile axis. However, the 0o/45o structure was strong along the 0o axis, and fracture was along the weaker 45o fiber interfaces, where the interlaminar sheer stress was high. Therefore, 0o/45o structures are expected to have a higher tensile strength than 90o/45o structures.

Therefore, 0o/45o structures are expected to have a higher tensile strength than 90o/45o structures. It is logical to select D-1 factor level for maximizing the tensile stress.Fig. 4. Fracture of Tensile Specimens345AC Interaction: The high temperature allowed a longer time for viscous flow of the ABS material to fill the open porosity. It also contributed to a greater degree of bonding between the fibers. Hence, it would be expected that the higher temperature (A+1) would increase the tensile strength. As previously explained, the low layer thickness (C-1) setting was preferred for increasing the tensile strength. Therefore, it is not surprising that the high temperature and low layer thickness (A+1C-1) would interact to maximize the tensile strength.Fig. 5. Trial 9 (Left) and Trial 11 (Right)Trial 9: Fracture surface of 0o/45o composite at low layer thickness.

5. Trial 9 (Left) and Trial 11 (Right)Trial 9: Fracture surface of 0o/45o composite at low layer thickness. Trial 11: Fracture surface of 0o/45o composite at high layer thickness.6.Conclusions Using an L16 (24) full factorial experimental design, the following conclusionsresulted from our work:1. The yield strength is maximized by low air gap, low layer thickness, low raster orientation and the interaction between high mold temperature and low layer thickness.2. The ultimate strength is maximized by low air gap and low layer thickness. 3.

The ultimate strength is maximized by low air gap and low layer thickness. 3. The effects of the factor levels on tensile strength were explained in terms of their microstructure and processing.Acknowledgment The work was funded by a National Science Foundation (NSF) grant to Loyola Marymount University for the Research Experience for Undergraduates (REU) program.References [1]Stratasys, FDM-1650 modeling machine, system documentation, web site: http://www.stratasys.com.[2]Es-Said, O., Noorani, R., Mendelson, M., Foyos, J., and Marloth, R. (2000), "Effect of Layer Orientation on Mechanical Properties of Rapid Prototyped Samples," Materials and Manufacturing Processes, 15 [1], 107-122.346[3] Montero, M., Roundy, S., Odell, D., Ahn, S-H, and Wright, P. (2001), "Material Characterization of Fused Deposition Modeling (FDM) ABS by Designed Experiments," Society of Manufacturing Engineers, Annual Conference, pp.

their interactions based on· the mechanical and physical properties of sintered ceramicspeci11lens. Post-processing software was developed to control and modify these parameters. [his software interface was .designed to mimic the QuicksHce™ interface for setting motion parameters based upon the material and the operation. Theresults of this investigation provided/useful information for the experilllentaLanalysis of the machine .specific process parfmeters.Suitable p~ratn~terswereselectedforthe EFF process fori fabricating representative ceralllicpr?totYl?es. \\Tiththe. optimized parameters, complicated parts were successfully fabricatedusil1g.bothKyocera SN282 and Starck M..11 silicon nitride powders.I. IntroductionExtrusion Freeform Fabrication (EFF) is an SFF technique based on the Stratasys FDMTM approacha for the fabrication of functional ceramic prototypes [1]. More details are given in reference 2.

More details are given in reference 2. While possessing the benefits of Fused Deposition of Ceramics (FDCTM) [1995 Volume from reference 1] and FDMTM,EFF has the added advantages that it can handle higher viscosity feed stock materials and higher extrusion temperatures compared to FDCTM and FDMTM. Similar to other SFF techniques, EFF also allows the sequential deposition of multiple layers to form a· complex ceramic shape. This has been achieved by. retrofitting a high-pressure extruder head to a Stratasys FDMTM modeler (figures 1 and 2). The CAD file is processed by the Quickslice™ software and used to control the EFFhigh-pressure extrusion head [T]. However, the operation parameters that are automatically set by Quickslice, are optimized for polymer filament type of feedstock.• These parameters are not necessarily the. optimum parameters for the SFF of ceramic parts.

optimum parameters for the SFF of ceramic parts. To further control process parameters for SFF of ceramic parts, SML Post™, a Visual Basic based post-processing software was developed to modify selected process parameters. SML Post™ could modify the start..delay,preflow, start-flow, start-distance, main flow , .\Vere •.•• originally set by roll-back, •• <speed, and •••• acceleration .•.• that the road-width, slice-thickness, and fill-patterns are still set by Quickslice™.shutoff-distance,Presently,a Stratasys Inc., Eden Prairie, MN103Quickslice. However, SML Post™ could modify these parameters also, if needed.An examination of the initial ceramic parts that were built using EFF suggested that furtheroptimization\Nasneeded to improve the part. quality. and surface .finish ofthe parts. This was the motivation for the work presented here. The optimization c0uldbeperformed on both machine and material parameters.

quality. and surface .finish ofthe parts. This was the motivation for the work presented here. The optimization c0uldbeperformed on both machine and material parameters. Material parameters include the type of ceramic powder used, the particle size distribution of the powder, the type of sintering aides used, and the type of binder system. Machine parameters are items such as.extrusion temperature, nozzle diameter, modeling envelope temperature, and percent flow (flow rate). For the work presented here, only machine specific parameters were optimized using the Taguchi method of Design of Experiments (DOE) [3, 4].Figure 1. Stratasys FDMTM retrofittedFigure 2. Close up view of the high pressureWith a high pressure headhead creating a ceramic prototypeII. Experimental ProcedureAn optimization of machine parameters was undertaken in order to improve part quality and surface finish.

Experimental ProcedureAn optimization of machine parameters was undertaken in order to improve part quality and surface finish. The part features used for optimization were dimensional accuracy, surface finish, and the time taken to build the sample. In order to reduce the number of experiments, the number of variables was reducedto seven. The same .sml file was used for all of the trials to reduce the number of variables. These variables could be changed using the post processing software (SML Post). The start distance and shutoff distance were held constant at values previously determined as optimal. These.numbers were obtained by trial and error. Earlier work had produced a binder system that demonstrated. favorable extrusion qualities when compounded with the ceramic powder [2]. The same. ceramic feed stock was used for the optimization experiments.

favorable extrusion qualities when compounded with the ceramic powder [2]. The same. ceramic feed stock was used for the optimization experiments. It is expected that minor modifications of the feedstock would result in similar optimized parameters for the EFF of those feedstocks.A. Optimization of the Perimeter OperationThe Taguchi method of Design of Experiments (DOE) was used for the statistical experimental design. Within the four types of operations used to deposit the material (perimeter,104contour, raster, and open), the parameters for perimeter were optimized first d~e to th~ relative ease and speed at which the experiments could be performed. AtwoJeveLexpenment wIth seven variables was performed to select and screen the most critical variables. The variables chosen for the screening experiment were start delay, pre-flow, start flow, main flow, rollback, speed, and acceleration.

The variables chosen for the screening experiment were start delay, pre-flow, start flow, main flow, rollback, speed, and acceleration. A description of the seven variables is given in table I.A Taguchi L12 array with three repetitions was chosen. With three repetitions, there was a 95% confidence level that a term identified as being significant truly belongs within the standard deviation as well as a 90% chance of finding a significant term if one actually existed. At this point, the predetermined high and low values of the different variables were entered into a computet program which created and analyzed DOE experiments (see table II). The software is called "Keep it simple statistically or KISS" supplied by Digital Computation. A three-level setofexperiInents was conducted to optimize the variables.Cylinders with a diameter of two inches and. a wall thickness of one road per layer were fabricated.

A three-level setofexperiInents was conducted to optimize the variables.Cylinders with a diameter of two inches and. a wall thickness of one road per layer were fabricated. The cylinders were built to a height of 0.2 inches.· An exit distance of 1.25 inches was used iniallcases. Twelve samples were made for each repetition.The parts· were analyzed in a number of ways. Thickness measurements from three points on the cylinders were taken for all 36 parts. The thickness before and after the starting point, and exactly opposite the starting point were measured. The standard deviation in part dimensions between the three points on each cylinder was calculated. The average thickness,the average excess.(an average of the measured thickness>minus .the thickness specified in the .sml file), time, and the end thickness minus the beginning thickness were also determined. By analyzing the statistics and graphs returned by the DOE software, the most important factors were found.

By analyzing the statistics and graphs returned by the DOE software, the most important factors were found. The DOE software analyzed the measured responses and calculated a factor known as "P(2 tail)". According to the DOE rule of thumb, if the "P(2 tail)" number for a particular variable is below 0.05, then that variable should be considered to belong in the regression model. If the "P(2 tail)" value was between 0.05 and 0.10, then that variable was marginally important. However, this variable would also be included in the three-level set of experiments. A variable with a higher value of "P(2 tail)" would be considered not to belong in the regression model.A Taguchi L9 matrix was created for the three-level experiments. The variables were chosen based on the screening experiments. The low and high values were retained from the screening experiment, while the middle value was an average of the low and high values.

The variables were chosen based on the screening experiments. The low and high values were retained from the screening experiment, while the middle value was an average of the low and high values. The parameters with a P(2 tail) number above 0.10 were set at values thought of as optimum, based on previous experimental results. The three-level Taguchi L9. matrix is shown in table III. The difference between the start and finish thickness of the parts created was measured and analyzed by the software. The optimized values of the machine parameters for obtaining the smallest difference in start and finish thickness were calculated by the DOE software. Several test parts were made with the optimized parameters.B. Optimization of the Raster OperationAs with the perimeter optimization, raster optimization was carried out using the Taguchi DOE techniques.

Optimization of the Raster OperationAs with the perimeter optimization, raster optimization was carried out using the Taguchi DOE techniques. The two-level experiment for the raster was done· with the same seven05variables as the perimeter two-level experiment plus the road width. The test shape for this set of experiments was a .1 "x2" rectangular plate with a rectangular hole .cut out of trle middle andi a uniform wall thickness of 1/8". Thirty-six (36) experiments were conducted. However, some\ofthese parts on a strictly qualitative scale of 0 to 10, 10 being the best. A road width of 0.0030" was used for all the experiments, since this closely matched the optimum road width obtained with the 0.0016" nozzle used.III. Results and DiscussionFOLoptimizingthe parameters for building the perimeter, the time to build the part and the difference in partthickness at the start and finish were used as responses and analyzed by the DOE software.

Since the buildtime yielded only one definite important factor (speed) and one possibly important factor (main flow), the difference in thickness at start and finish was usedin constructing the three-level experiment. •Theeffectof the different parameters on the difference in part thickness at the start and finish is shown in figure 3. The slope of the marginal means plot in figure 3 is indicative of the significance of the different build parameters. From figure 3,it was se.enthat the start-delay, main-flow, rollback, and speed were significant factors for the optimization of the perimeter. The Taguchi L9 matrix created for the three-level· experiments based on •these> results was •shown in·. table III. The difference between the start and finish thickness of the parts. created was measured and analyzed by the software.

table III. The difference between the start and finish thickness of the parts. created was measured and analyzed by the software. The program returned values for the P(2 tail) which confirmed that all of the variables· that· were used at this experimental stageiwere significant in minimizing the difference between the start and finish thickness. \The·optirriizedvalues of the.machine parameters for obtaining the smallest difference in start and finish thickness are shown in table IV. These values<. are for a particular powder/binder system and tip size. Further optimizations will have to be performed in order to obtain the same kind of results for different tip sizes and different feedstock material.

Further optimizations will have to be performed in order to obtain the same kind of results for different tip sizes and different feedstock material. In the case of raster, the start-delay, main-flow, roll back,speed and road width parameters were observed to be the most significant.It was foundthat when usingJhe optimized parameters, the surface finish of the parts was greatly improved with respect to parts that were built previously without optimization. Fourtest cylinders were made using the optimized parameters for perimeter. The most dramatic change occurred with the build speed. The build speed suggested by the optimization program was 0.3" per second, compared to 0.8" per second before optimization. The surface finish and the seam where the layer was started and finished showed no pooling of excess material. Figures 4a and 4b shows cylinders that were made with minor changes in the parameters, while figure 4cshows a cylinder that was built with the optimized parameters.

Figures 4a and 4b shows cylinders that were made with minor changes in the parameters, while figure 4cshows a cylinder that was built with the optimized parameters. Sample parts with good surface finish and part quality was built using the optimized raster parameters. As indicated above,optimum part quality was obtained with a road width of 0.030" when using a tip size of 0.016". Further optimization experiments are in progress with tip sizes of0.012" and 0.025".Mechanical testing was completed for Starck silicon nitride samples fabricated with the optimized conditions.· Prior tests had \shown that post-polishing provided increases in strength. Samples were made with both 0 degree and alternating 0/90 raster directions. The strength values for all test bars can be seen in table V. For samples with a 0 degree raster build, the strength datawas similar for polished and unpolished specimens.

The strength values for all test bars can be seen in table V. For samples with a 0 degree raster build, the strength datawas similar for polished and unpolished specimens. The test bars that were made with 0 and 90 degree alternating raster angles showed poor surface finish, compared to 0 degree only, and thus the data shows a difference in strength values for polished and unpolished samples. Because of the geometry of the test bar, a 90 degree raster angle is a poor choice since the road would only be about 0.1 inches long. The smaller the road, the more chances of introducing a defect into thepart.Other silicon nitride powders have been used for EFF with the same binder system and extrusion conditions with favorable results. Two sample parts fabricated using the optimized conditions are shown in figures 5 and 6. In figure 5, green parts made with Kyocera SN282 silicon nitride and Starck M-l1 silicon nitride powder are shown.

In figure 5, green parts made with Kyocera SN282 silicon nitride and Starck M-l1 silicon nitride powder are shown. Figure 6 shows a turbine blade that was made with the optimized parameters.IV. ConclusionThis investigation has established the feasibility of using DOE techniques for the optimization of the. EFF process when creating functional ceramic prototypes. For building the perimeter, the parameters that control the EFF process quality are start-delay, main flow, speed the start-delay, main-flow, roll back, speed and road width and rollback. For the raster, parameters were observed to be the most significant. It has been demonstrated here that with the optimized parameters, the overall part quality was improved. Complicated ceramic shapes were built by the EFF process using the optimized parameters. Currently more optimizations are planned forother tip sizes.V. AcknowledgementsThis work has been performed under NASA STTR program contract number NAS8 98025.

Currently more optimizations are planned forother tip sizes.V. AcknowledgementsThis work has been performed under NASA STTR program contract number NAS8 98025. The authors would liketo thank Jiong Peng of the University of Arizona for her help in the preparation and optimization of the EFF samples.References1. Proceedings of the Solid Freeform Fabrication Symposium, Vol. 1990-1998, Edited by H.L. Marcus, et aI., the University of Texas at Austin, Austin, TX.2. R. Vaidyanathan, J. L. Lombardi, B. Tennison, S. Kasichainula and P. Calvert, "Rheological properties of ceramic formulations for Extrusion Freeform Fabrication," presented at the 23rd Annual Cocoa Beach Ceramic Society meeting on Ceramics and Composites, Cocoa Beach, FL, Jan. 1999 (in press).3. Ross, Phillip 1., Taguchi Techniques for Quality Engineering, McGraw-Hill: New York NY . ,4.

,4. Schmidt, Stephen R., Launsby, Robert G., Understanding Industrial Designed Experiments ' "Extrusion Freeform Fabrication of silicon nitride parts,"Proceedings of the 10th Solid Freeform Fabrication Symposium, (1999), in print.107Table I: A description of the software controlled (process) parameters.Para.meter Start delay Pre-flow Start flow Startdistal1ce Main flow Shutoff distance Rollback Speed AccelerationDescription The lag time between layers Extrusion rate during stops Extrusion rate at the start of the exit distance Distance from start point for main flow to start Extrusion rate over the bulk of the layer The distance from the stoppointfor main flow to stop The rate the material is pulled backto reduce the flow Therate at which the head moves along a set path The rate at whiCh the extrusion head changes speedTabl: e ! F#lr.tnr Rnw# Start 0II Th Te Ad'h' L12 a~ uc 1 Bh matrlxuse mt e screenmg expenment or penmeter optlmlzatl D E C FlolJl.

Cylinder c was constructedwith the optimized parameters.Figure 5: Impellers made with Kyocera andFigure 6: A turbine blade madeStarck silicon nitride powders.with the new parameters.110.

plication of Factorial Design in Selective Laser SinteringXiaoming (Michael) Deng and Joseph J. Beaman Department of Mechanical Engineering The University of Texas at Austin, Austin, TX 78712AbstractSelective Laser Sintering (SLS) is a complex process involving many process parameters. These parameters are not all independent. A factorial design technique is utilized to study the effects of three main process parameters, laser power, laser beam scanning speed, and powder packing density as well as their interactions on the sintering depth and fractional density. The results of this investigation provide useful information for the further experimental analysis of the process parameters and for selecting suitable parameters for SLS process.1. IntroductionThe SLS processing parameters and their interactions have strong effects on the quality of SLS products. These parameters are summarized in Figure 1.

IntroductionThe SLS processing parameters and their interactions have strong effects on the quality of SLS products. These parameters are summarized in Figure 1. An experimental cross parametric evaluation and a parametric analysis effort has been conducted to assess the effects of the major process parameters on the quality of the SLS products [1]. It was focused on the investigation of geometric phenomenon, sectional sintering geometry, and density in a single layer basis. For simplification, the effects of environmental influences and material were fixed. In other words these factors were treated as constants. Machinery influences were neglected other than the powder packing density. Many interesting results of the effect of the individual processing parameter on the quality of the SLS products have been achieved in that experimental research. However, the effects of the interaction of these parameters on the quality of the SLS products are still unclear.

However, the effects of the interaction of these parameters on the quality of the SLS products are still unclear. In this paper a fractional design technique was utilized to to determine the most effective factors for SLS part production and to study the interactions of the parameters. Sintering quality is described using sintering depth and fractional density and is examined in terms of major input parameters including laser power, laser beam scan speed, and powder packing density .2. Experimental DataThe experimental approaches used in this study are described in Reference [1] in this proceedings. The sintering depth and fractional density data for different laser powers, scan speeds, and powder packing densities are listed in Table 1 and Table 2, respectively. Sintering154depth can be identified from a so-called sectional sintering geometry. Sectional sintering geometry is a section view of the sintered part (path).

Sintering154depth can be identified from a so-called sectional sintering geometry. Sectional sintering geometry is a section view of the sintered part (path). It is defined from sintered surface, alongFixed for this research........ Laser system _Scanning--- systemI Beam spot size -1 Laser power H Wavelength Y Beam profile rl Scanning speed r-H Scanning path I I Scanning distance I (geometry)ProcessParametersSelective Laser SinteringfelllllIIFigure 1 Parameters in SLS processthe direction of laser beam (Z-axis), to a region that sintering density appears obviously reduced (Figure 2). The sintering depth is defined within a solid sintering region. It is155I I I Ievaluated by means of the mean heights of the solid sectional sintered portions. The sectional sintering geometry can be clearly observed and measured using SEM.

It is obvious that the pressed density, laser power and scanning speed effect sintering depth. The effective degree of these three factors to the sintering depth is ordered from high to low: pressed density, laser power and scanning speed. The scanning speed-laser power, laser power-pressed density and laser power-pressed density interaction also affect the sintering depth. The scanning speed laser power-pressed density interaction is significant at 5 percent, indicating a relative mild interaction among these factors.Table 3 Analysis of Variance for sintering depthSource of VariationSum of SquaresDegrees of FreedomMean SquareFoScanning speed (A) Laser power (B) Pressed density (C)AB AC BC ABC Error Total0.052206 0.062082 0.422256 0.025730 0.013883 0.010447 0.005679 0.012409 0.6046933 1 1 3 3 1 3 48 630.017402 0.062082 0.422256 0.008577 0.004628 0.010447 0.001893 0.00025967.32 a 240.15 a 1633.41 a 33.18 a 17.90 a 40.41 a 7.32 ba Significant at 1 percent.

b Significant at 5 percentThe analysis of variance for fractional density is summarized in Table 4. In this calculation, the results show that fractional densities are dominant by laser power. All of the factors and their interactions are significant at 1 percent. The interaction of the laser power with the packing density are the second largest effective factors to fractional densities.3.3 Experimental errorsThe major experimental errors come from the stability of the laser power and the environmental temperature. But the error can be minimized if the laser is at steady state before every experiment start. The calibration of the laser power for each setting is necessary. In158general, the experimental errors can be greatly reduced if the experiment is properly conducted and the initial conditions are properly controlled.

In general, measurement error of the SEM is less than 10%, and this error can be controlled if the SEM is calibrated before each measurement.Sample preparation can also cause a measurement error if the viewing section is not paralleled to the surface of the SEM sample base. However, this error can be reduced by sample preparation since the titled angle is generally smaller than 2 to 3 degrees, and the projection effect of this angle is too small to affect the measurement in normal circumstance.3.4 DiscussionsAs the results of the factorial design, the three selected factors: laser power, scanning speed and packing density are significant to sintering depth and fractional densities.

The results show that1) the sintering depth is dominant by the powder pressed density,2) the sintering fractional density is dominant by the laser power,3) laser power effects the sintering depth more than scanning speed does,4) scanning speed affects fractional density more than laser power,1595) the interaction of the laser power and packing density is significant to the fractionaldensity,6) the interaction of three factors is significant to the sintering depth only at 5%.The experimental and measurement errors are small. In this investigation, the research interest is focused on a parametric qualitative analysis that are most important SLS processing.40 Summary and ConclusionsThe results from factorial design indicate that these three factors are all significant at 1% to the sintering depth and sintering fractional density. Their interactions are significant to sectional sintering geometry as well.

Their interactions are significant to sectional sintering geometry as well. Sintering depth is affected most by the powder pressed density. Sintering fractional density is affected most by the laser power. Further investigation should be conducted more thoroughly to study the effects of all the factors significant to the SLS process on the SLS part quality.AcknowledgementsWe acknowledge Dr. Willis Adcock for his helpful comments in this research.References1. X. Deng, G. Zong, and J. J. Beaman, "Parametric Analysis for Selective Laser Sintering of a Sample Polymer System," Proceedinis of Solid Freeform Fabrication Symposium, edited by H. L. Marcus, J. J. Beaman, J. W.. Barlow, D. L. Bourell, and R. Crawford, Austin, TX, August 3-5, 1992.2. D. C. Montgomery, Design and Analysis of Experiments, John Wiley & Sons, 1991.160.

plicationthe ..;'....:",.....o..J/'~,'"'d .....Jlu. of Multi-Material StructuresManufactured in 'llr..UII"'.I.'-lIl PrototypingOliver Konig <okoenig@ikb.mavt.ethz.ch> Institute for Construction and DesignZurichCH-8092 ZurichGeorges M. Fadel <gfadel@ces.clemson.edu> Department of Mechanical Engineering Clemson University Clemson, SC 29634-0921, USARecent developments in the Rapid Prototyping technology establish it as a new manufacturing technique, enabling localized material addition to build a part. Thus, heterogeneous structures, consisting of more than one material can be produced. The aim of this paper is to present an optimization tool to find the best material distribution in a multi-material structure due to given objectives and constraints. The tool is based on genetic algorithms using a discrete material model and FE"'analysis to evaluate the objective functions.

The tool is based on genetic algorithms using a discrete material model and FE"'analysis to evaluate the objective functions. It can optimize the distribution materials in 2D-structures with up to 1500 DOF's at reasonable VVA,Jl.Jl.i-/ ... "'","'AV.LA.....'" costs. Its performance is shown on a bi-objective optimization of a turbine blade.1the past, the intuition and experience of engineers played the key role in designing seen the development of numerical tools, which provide conceptual aim of these tools is to experience of an engineer. addition, most of these tools are traditional structures consisting of one material. Examples of such tools are the topology optimization method using homogenization introduced by Kikuchi [1], and the 'Soft-Kill-Option' (SKO) method introduced by Mattheck [2].

Examples of such tools are the topology optimization method using homogenization introduced by Kikuchi [1], and the 'Soft-Kill-Option' (SKO) method introduced by Mattheck [2]. development of new manufacturing technologies and new materials such as Rapid Prototyping [3] or the use of composite structures expands the demands for numerical tools to possibility to manufacture anisotropic and/or design structures. These new techniques offer multi-material structures. This makes it to design parts just by intuition. the Numerical optimization algorithms are needed to develop solutions which consider all different aspects of such a problem. The present project starts at this point, concentrating on the design of multi-material structures manufactured in Rapid Prototyping. This leads to the main research question for this paper:structures. designsyearsa given design space and specified boundary conditions.

This leads to the main research question for this paper:structures. designsyearsa given design space and specified boundary conditions. intuition andfocused on the optimizationmoreHow can one optimize the distribution ofdifferent materials in a multi-material structure for given objectives and given constraints?The paper presents a methodology based on genetic algorithms to solve this type ofoptimization problems.2 Genetic AlgorithmsGenetic Algorithms can be described as search algorithms based on the mechanics of of stochastic search methods,natural selection and natural genetics. They belong to a209notby ......... u'''''v ...... \,......J.uv'~u .....,'"'uJ.j;;:,J.J. st)ac(~s.............' .............'J....J.'-I'\,...""4.....J. programming principles. On the other side there requirement associated with the use good solution approach techniques are inefficient. • • GA's proceed from several points in the design space,this method.

On the other side there requirement associated with the use good solution approach techniques are inefficient. • • GA's proceed from several points in the design space,this method. Therefore genetic ........ F\v .........J.J.J. ... J.u n:mn~Se]ltstandard mathematical can be formulated as follows:work on function evaluations,to find globaloptima.GA's work on a coding of the design variables.allows them to workdesign spacesconsisting of a mix of continuous, discrete, and integer variables.GA's in Automated 1l8llJ'·U~1DMuch workUUl,IJ.J.i:lJ..l....U as shown below:u..lL"'",u.L"L"V.:"l ~""J.J."".L~"".L l,.l..:HJ....'v"..:J ofgenetic CP'!:l1"l"hand topology optimization. structural topology design. Nair concepts with genetic optimization with a genetic University.Jakiela [6] apply genetic algorithms to problems of [7] investigate the combination of approximation flywheel structural optimization procedures.

presented by the GARAGe .[8] group at Michigan StateOther projects are more concerned in the application of genetic search on practical problems. For example there are papers investigating multidisciplinary rotor blade design [9], the design of a satellite boom [10] or the optimization of truss structures [11].Finally, several publications examine the optimization of composite structures withgenetic algorithms [12],[13],[14] . All these works establish the use of genetic search in automated design optimization.LonC4ent and AdditionalGA'sThe concept is briefly discussed in this paragraph, a detailed description canfound inthe book frol11 Goldberg [4].After the coding of the design variables an initial population is generated randomly. Then fitness values for each individual in a population are evaluated. the iteration process starts. The first GA operator which is applied to the population is reproduction.

Then fitness values for each individual in a population are evaluated. the iteration process starts. The first GA operator which is applied to the population is reproduction. Individuals are selected for the next generation according to their fitness. The crossover operator mates and crosses the individuals in this newly generated pool ofindividuals. Mutation as the last operator v"' ..' .....UJ.v............. random alteration of the value of a string position. This concludes one iterationnew generation results.some additional"GA features which are used in the project are briefly introduced:Elitist • Overlapping Populations. The pool of individuals before reproduction consists of a GA with overlapping populations in the previous population and a specific amount of new individuals. pool are removed in order to return the population to its worst individuals of theThis strategy ensures, that the best individuals stay in the population.original size.

pool are removed in order to return the population to its worst individuals of theThis strategy ensures, that the best individuals stay in the population.original size. Since only part of the population is generated, this strategy saves computation time. Scaling the Objective Function. In order to control the sensitivity of the GA, the objective values are linearly scaled. The aim of a scaling function is to avoid premature convergence in the GA.2.3 The initiation ofgenetic search requires specification ofsome key parameters: •Parameters in Genetic SearchPopulation Size:. popsize. The number of strings processed in each generation must be kept small to· minimize the overall computation· effort.

popsize. The number of strings processed in each generation must be kept small to· minimize the overall computation· effort. A population size between 25 and 125 represents a good choice for structural optimization problems [5].Number of Generations: ngen• Usually a value in the hundreds is needed to make sure that the solution has time to converge.CrossoverProbability:pcross.Valuesranging from 0.6 to 0.8 have been used in numerical experiments with very)satisfactoryr~sults.Mutation Probability:Pmut.Probability values between 0.005 and 0.05 produce in general good results.OverhlJ.>ping Gap: Prepl. The overlap parameter specifies how many new individuals are created for each generation. A typical value is Prepl=0.5, meaning that 50 % of the population has to be evaluated new for each generation.3 Modeling of Materials and Structures3.1 Discretization ofthe Design DomainAny kind of chromosome used as a representation of the design variables in GA's isdiscrete.

Therefore a finite design domain must be selected and discretizedinto elements.For atopology optimization the design domain represents the maximal volume in which the structure is to be constructed. For a multi-material optimization, one can either fix the the topology and material topology and only optimize the material distribution, or both, distribution can be unknown.The representation of the design space in a mathematical way is the space T = R 3 [15]. A typical obJectin the design space is called a class A, while members of this class are noted as r-sets, for example an element of the discretized structure.3.2 Modeling of Multi-Material StructuresCurrently, most solid modeling techniques are capable of capturing only the geometric information of an object.

In this project, a modeling method for generaland topological heterogeneous objects proposed by Kumar and Dutta [151 is used:Modeling of Heterogeneous Objects ConSisting of a Finite Number of .• Distinct Materials (HD)116]. In the previous paragraph, the design space T = R 3 was definedin which t1}egeometry of an.obJectcat1l'erepresented. In order to model multi-material objects,this spaceofintegers I .• TheproductspaceT =R 3 xI formsthentl1enewmodeling§pac.e including both, geometrical topology and material distribution. A new class Am=A x K is defined where K c is a finite set of integers. A typical member of this class is then defined as anrm-set. It is211composed of an r-set describing the geometry and an integer k E K defining the material. Each rm-set, typically an element in this project, is homogeneously filled with one material.The design variables for the optimization are now defined with the type of material for each element in the discretized design space.

The approach provides thus a discrete and finite set of design variables.3.3 Mapping the Design Variables to GA ChromosomesThe design vat'iables for.a GA1111.lstb~coded,••• amappingdomain to chromosomes must be defined. Using the HD approach above, the coding results a binary or integer string which defines the type of material for every element inthe design domain. A binary string .can be used if there are only two materials to distribute, while an integer string is able to handle a whole set of materials.a cantilever structure with three materials mappe<i into a two dimensional integer string. To consider topology optimization as well as optimal material distribution, the material 1 is defined as void. To prevent numerical singularities, the material void is implemented with very small property values.Figure1 showsthe mapping.

To prevent numerical singularities, the material void is implemented with very small property values.Figure1 showsthe mapping. for223111111223331111222222222222222222223331111223111111Figure 1: Mapping to GA Chromosome4 Optimization ProcedureThe genetic algorithm library GAlib in C++ by Matthew Wall (MIT) provides all theneededfunctionalityand is used in this project.A GA needs a function which determine the fitness of each individual in a population. For a multi-material. optimiJ!';ationthisfunctionincludes an analysis of the discretized structure. For a typical GAevaluation, the fitness has to be computed for thousands of individuals. Therefore it.is imperative that the analysis method.has to be fast. This can only be reached if the method for the fitness evaluation has a close interface to the GA.A first . approach to evaluate the fitness of the individuals is to use Numerical Approximation Concepts.

approach to evaluate the fitness of the individuals is to use Numerical Approximation Concepts. Keane [7] used this concept to optimize a lObar truss structure. It is based on the exact analysis Jor a limited number of individuals, the fitness for the rest of the population is evaluated using an approximation model. Although this approach can save a lot of computation time, one would have to verify the accuracy of the results carefully, because only a small part of the individuals are evaluated exactly. This method was not pursued any further in this project.in the genetic algorithm represents another approach to evaluate the fitness values required by the genetic search. This method is favored for this project, because it provides in general a good accuracy for the solution and adaptionsofthe method can be found for analyses in a lot of different fields. The computational cost for analyzing each individual in the GA is the main concern.

The computational cost for analyzing each individual in the GA is the main concern. Therefore a big effort was put in minimizing the computational time needed for one single analysis. In order to keep the Qptimization tool as compact as possible, a FE-code in C++ was developed. With this approach, the optimization procedure can be packed in one single C++ program, the time consuming processes such as storing temporary data on the hard drive are eliminated.A Finite Element Analysis for each individualThe FE-code was developed to provide analyses for the applications in section 5. Therefore FE-codes for heat conduction and thermal stress analyses had. to.be implemented. The concepts of the FE-codes can be found iu[17],[18],[19], and are not discussed here. The detailed procedure is. shown in the flow chart in Figure 2.The computational performance of this app roach was tested for a simple topology optimization of a· Cantilever.

The detailed procedure is. shown in the flow chart in Figure 2.The computational performance of this app roach was tested for a simple topology optimization of a· Cantilever. It represents a small application with 352 DOFfs, where not much.computation time should be needed to solve it. In a typical optimi zation for this problem the fitness function has to be evaluated The evaluation on a computational server with 2 CPU's (200 MHz) and 1 GB RAM needs approximately 75s. This is promising for bigger applications.approximately3000times.In an earlier approach Ansys was used for the FE part. But the time consuming processes of storing interface files on the hard disk makes the GA evaluation for the same application about 300 times slower.,IIIStart GA 1 Read in Constant FE-DataIIIInitialiZe Start population: n='..Population nII1 lndl-,Vld_ua_II__----'1 ..

I Transfer String to Element Materials IrlL-I)"IS 1 I I Write BestIConvergence Crtteria 11--_-, Isatisfied?1noII IIndividuals to File IReproductionICrossover.. • I Population n=n .., IIMutationII IThe optimization procedure using the FEII Terminate GAilCode in C++ is therefore used for all applications.Figure 2: Procedure Using FE-Code in C++5 Multi-Material Optimization of a Turbine BladeThe· application in this section shows the performance of the developed optimizationmethod. The distribution of two materials in a turbine blade shall be optimized.5.1Problem DefinitionThe design domain considered is a 2D cross section of a turbine blade. It is a non-cooled blade and has therefore no internal holes in the structure. This cross section is discretized into 951 linear triangle elements.Materials. A turbine blade is subject to pressure gradients over the boundary, which introduce high mechanical stresses in the structure.

A turbine blade is subject to pressure gradients over the boundary, which introduce high mechanical stresses in the structure. Furthermore, certain regions of the blade in a gas turbine are heated to very high temperatures. This results in the wish to use one material, such as Titanium, to deal with the mechanical stresses in the structure, and a heat withstanding material such as a Ceramic to place at high temperature locations. Therefore, in this example, the materials Titanium Alloy and Silicon Nitride are used.Boundary Conditions. The temperatures applied on the boundary are based on experimental data described in [20]. Figure 3 shows the resulting temperature distribution for a titanium blade. The purpose of this application is to show the optimization algorithm. Therefore the complexboun.dary conditions on a real turbine blade were simplified: • • The bladeis not cooled, there are 110 in.ternal holes in the structure.

Therefore the complexboun.dary conditions on a real turbine blade were simplified: • • The bladeis not cooled, there are 110 in.ternal holes in the structure. •Prescribed temperatures on the boundary replace convection and radiation effects.Pressure distributions on the boundary are not applied.2135.2 The objective is tp filld the optimal material service to given ..ulaximal due distribution an for temperatures is the materials. This unconstrained optimization problem. The fitness function for the genetic algorithm is formulated as a maximization problem:Single Objective OptimizationMaximizescore = Nelescorenorm1200K975 1000 1025 1050 1075 1100 1125 1150 1175 1200"2 L..J (1; - T:ev/ce(mat») /=1Figure 3: Temperatures.inaTi Blade[K] The service temperature ~erv/ce(mat) does not represent a physical property.ofthe material,it is fictive.

Since the conductivities for the two materials are different, the temperature distribution in the blade changes for every new chromosome. A heat conduction analysis has to be performed for each individual in. the GA. The design variables are represented by the material of each element, they are mapped in a binary$tring fpr the GA. The GA.is evaluated for 2000 generations with a population size of 100. This corre$ponds to 100'000 heat conduction analyses. A typical evaluation takes about 70niin. Figure 4 shows the fitness scores for this evaluation, the best individual found is shown in Figure 5.1300r---~----~---,1200 .1100 .1000I 900 . iI 800 ... .e u.. 700 .......... ; .......j,Titanium Alloy...................................... ..

scorestressr1Where Wtempand wstressrepresent the weighting factors. The objective functions are defined as:NeleL (1'; - Tsev/ce(mat})2NeleL O'eqSCOretemp =/=1nOrmtempSCOrestress =..-!.:./=~I__nOrmstress214The temperature objective is defined similarly to section 5.2, it controls the material distribution according to temperature distribution. The thermal stress objective is aims at reducing the thermal stresses in the blade, it is the sum of an equivalent stress O"eq of allelements. To evaluate the fitness of a single individual, a thermal conduction and a thermal stress analysis have to be performed.In a bi-objective optimization, many optimal solutions can be found. Therefore a Pareto set [21] of solutions is created in a first stage. The Pareto points result from a systematic variation of the weight factors between 0 and 1. Figure 6 shows the resulting Pareto points for such a variation.

The Pareto points result from a systematic variation of the weight factors between 0 and 1. Figure 6 shows the resulting Pareto points for such a variation. Each of this points represents the best population found by the GA in 1500 generations with a population size of 100 and the specified weight factors. The temperature objective is plotted in the x ..direction, the stress objective in the y-direction. A typical GA evaluation for this problem takes about lOh, because of the two FE-analyses (thermal conduction & thermal stress) for each fitness evalUation.A decision making process is necessary to choose a single solution out of this set of optimal Pareto solutions[21]. Figure 7 shows one single solution where the weight factors were chosen as wtemp =0.3, Wstress =0.7. In addition the GA for this result is run over 2000 generations with a population size of 150 to get a more converged result.,.jNt=0.9 ws=0.:1 0.5 ~';O:8 'WS;,;O:? ·: : .".

In addition the GA for this result is run over 2000 generations with a population size of 150 to get a more converged result.,.jNt=0.9 ws=0.:1 0.5 ~';O:8 'WS;,;O:? ·: : ." . ..IO.49:ewt=0.V·ws",;O;3J:eM=0.5 w$o:O.50.48 ..~.":();6,,!,s:,,:p'.:4.. ";."wt=0.1 WS=lj9J;Titanium Alloy:eM=9.3 ws=0.7jSilcon Nitride0.47 ..:wt=O.2·ws=O;S;,.:lIWt=0.4 w~=0.6·: . .0.46 ' - ' - - _ - - '0 -__--'-__----'_----10.350.40.450.5ItempHFigure 6: Pareto PointsFigure 7: Best Individual for wt = 0.3,w t =0.3emps ress6 Discussion and ConclusionFor the single objective optimization in section 5.2, the optimization puts as expected Ceramic at the tip of the blade (see Figure 5), where the highest temperatures occur. The Titanium in the rest of the structure provides the needed strength. Although the GA was evaluated for 2000 generations, it's stochastic concept still influences the result.

The Titanium in the rest of the structure provides the needed strength. Although the GA was evaluated for 2000 generations, it's stochastic concept still influences the result. Several elements have obviously the wrong material (assigned by mutation). Because of the high temperatures and the different thermal expansion coefficients, it is imperative to include thermal stresses in the optimization.An approach to include thermal stresses is shown in the bi-objective optimization in section 5.3. Figure 6 shows that the weight variation does not result in a convex pareto curve. This indicates that probably the GA evaluations did not fully converge to the optima.215Figure 7 with the best individual for a specific weight combination shows that the Ceramic is smaller. But it is still not clear how the interface between the two materials should be best configured.

But it is still not clear how the interface between the two materials should be best configured. It was hoped to see a fuzzier mechanical interface between the two materials, but the problem conditions and the way the objectives were formulated prevented this from happening.An optimization tool for 2D-multi-material structures was developed. It optimizes the distribution of materials in a structure due to given objectives and constraints using genetic algorithms. It can be said, that the developed optimization tool is able to handle structures with as many as 1000 to 1500 DOF's in a reasonable computation time (a few hours). The turbine blade application showed, how essential a fast objective function evaluation for GA's is, a run time of 10h for the bi-objective optimization shows the limits of the tool.

The turbine blade application showed, how essential a fast objective function evaluation for GA's is, a run time of 10h for the bi-objective optimization shows the limits of the tool. Since the solution space for the GA grows exponentially with the number of elements, the number of generations needed to achieve a converged solution increases as well in an exponential way. This project has shown some of the potentials of GA's in structural optimization, as their flexibility, their robustness and ability to find global optima. But it has also demonstrated some of their shortcomings of high computational needs and bad convergence performance towards the end of an optimization.

But it has also demonstrated some of their shortcomings of high computational needs and bad convergence performance towards the end of an optimization. To overcome these shortcomings, future work could be concentrated on introducing parallel genetic algorithms, using gradient based optimizers after the GAs pointed towards possible solutions (memetic algorithms) or using a continuous material model.References[1] Bends0e, Martin Phillip,Optimization ofStructural Topology, Shape, and Material,Springer-Verlag Berlin, Heidelberg, New York, 1995[2] Konig, 0., Application and Analysis()fa Topology Optimization Method Including the Development ofa Bicycle Crank, Semester Work at the ETH Zurich, in German, 1997[3] Beaman, 1.; Barlow, J.

PLICATION OF LASER-BASED ADDITIVE MANUFACTURING TO PRODUCTION OF TOOLS FOR FRICTION STIR WELDINGJ.H. Ouyang, H. Mei, M. Valant, and R. KovacevicResearch Center for Advanced Manufacturing, Dept. of Mechanical Engineering, Southern Methodist University, 1500 International Parkway, Suite #100, Richardson, TX 75081, USAAbstractlaser-based additive manufacturing to production of tools for friction-stir welding (FSW) of materials. The design and rapid manufacturing of powerful tools has become one of the major challenges in expanding the application of FSW processes to new materials. The one-step laser cladding process is capable of creating a novel, fully dense and metallurgically bonded near-net-shape tool with microstructural and compositional gradients. In this paper, two kinds of FSW-tool materials are developed and evaluated for the applications of welding different materials.

In this paper, two kinds of FSW-tool materials are developed and evaluated for the applications of welding different materials. A pure H13 tool steel powder is used to build a FSW-tool for welding the materials with relatively low melting temperatures, such as aluminum alloys. A WC-based ceramet/tool steel functionally graded material (FGM) is synthesized to build a FSW-tool in order to weld the materials with higher melting temperatures and highly abrasive materials such as MMC materials. By controlling the amount of different supplied powders under the optimized laser cladding conditions, WC-based ceramet/tool steel FGMs are successfully synthesized layer by layer.

By controlling the amount of different supplied powders under the optimized laser cladding conditions, WC-based ceramet/tool steel FGMs are successfully synthesized layer by layer. FSW experiments are carried out to evaluate the deposited tools.This paper presents a feasibility study of application ofIntroductionA novel solid-state joining process, friction stir welding (FSW), invented by The Welding Institute (TWI, Cambridge, UK) in 1991, has recently caught the great attention of the welding community for fabricating high-quality butt and lap joints of aluminum alloys. The same potential exists for FSW of hard metals such as titanium, stainless steel, and nickel base superalloys [1-3]. One of the major challenges in expanding the application of FSW processes to new materials is the lack of suitable tools for welding materials with high melting temperatures [1].

One of the major challenges in expanding the application of FSW processes to new materials is the lack of suitable tools for welding materials with high melting temperatures [1]. To be effective, tool materials must resist physical and chemical wear, possess sufficient mechanical strength at elevated temperatures, and effectively dissipate the heat carried to the tool during the welding process. The development of new tooling made of polycrystalline cubic boron nitride (PCBN) appears capable of meeting the requirements, but the cracking of tool during the FSW of materials clearly causes the degradation of the tool life [1]. The suitability of FSW for hard metals such as stainless steel and titanium alloys has been investigating with a carbide tool [1-3].Most manufacturers have been looking for a device that can make a tool directly from a computer-aided design (CAD) drawing with a designed composition, microstructure, and properties.

Selective laser sintering (SLS), laser-engineered net-shaping (LENS), micro-casting, laser-based additive manufacturing (three-dimensional laser cladding), and shape deposition65layer-manufacturing processes for the rapid manufacturing (SDM) are some important prototyping/manufacturing system. In the one-step laser cladding, the powder material is injected continuously into a laser molten pool and melts to form a clad track on the surface of the substrate. A layer is formed successively by tracks deposited side by side [4,5]. A variety of materials including metals, ceramets, ceramics, and composite claddings have been successfully deposited to develop a potential in non-equilibrium synthesis of advanced functional materials, alloy development and free-form near- net shape manufacturing [4,5]. Composite surface layers have been produced by injection either ceramic particles or a mixture of metal and ceramic powders at a certain ratio, into the laser molten pool [4,5].

Composite surface layers have been produced by injection either ceramic particles or a mixture of metal and ceramic powders at a certain ratio, into the laser molten pool [4,5]. Hard particles such as WC, TiC, SiC, and CrB2 are usually employed to improve the wear resistance of the engineered surfaces [4,5]. WC particles are distinguished by a minimal plastic deformation capacity, a low thermal expansion and a high wettability by molten metals. With a high content of hard particle, the deposited surfaces are relatively rough and a high crack rate cannot be avoided. This result may be overcome with the use of solid state Nd:YAG lasers that allow improved control of the cladding process at low heat input [4]. High-power lasers in conjunction with a powder- feeding technique, have been used to produce functionally graded materials (FGMs) by the successive deposition of different clad layers.

High-power lasers in conjunction with a powder- feeding technique, have been used to produce functionally graded materials (FGMs) by the successive deposition of different clad layers. However, the large differences in thermal and physical properties between metals and ceramics cause serious problems in the fabrication of metal/ceramic FGMs by laser cladding [4]. Since the heating and the cooling rates are rapid during laser cladding, the cracks form more easily not only at the interfacial regions between different constituents but also within the ceramic particles. There are difficulties in controlling the porosity and cracking in some cases, and the chemical interaction between the matrix and ceramic particles due to high temperatures involved in laser cladding.In this paper, a feasibility study has been carried out to produce a FGM FSW-tool consisting of WC-based ceramet and H13 tool steel using the laser-based additive manufacturing technique.

A graded distribution of WC particles was realized by adjusting the flow rate of different powders for successive layers. The WC particles in the deposited FGM provides a high wear resistance, while the H13 tool steel exhibits an excellent combination of hardness and toughness at elevated temperatures.Experimental proceduresFig. 1 shows the schematic diagram of the laser-based additive manufacturing system setup used in this investigation. A 1kW continuous wave Nd:YAG laser with a 200- mm focusing optics is used to deposit the powders. Recently developed computer-controlled powder feeders are used for dosing and feeding the powder mixture into the desired composition, and allow for an exact setting and continuous change of the powder mixture during the cladding process. The alignment of the laser beam, powder nozzles, and their positioning in relation to the substrate must be accurate for the process to be reproducible.

The alignment of the laser beam, powder nozzles, and their positioning in relation to the substrate must be accurate for the process to be reproducible. The feed rate of each of the powders and laser parameters are controlled automatically. All the deposited parts are built up on the substrate of an AISI 1018 mild steel. The clad materials studied are H13 tool steel powder and WC-based ceramet powder. The chemistry of H13 tool steel powder with a particle size range of +44 to -152 µm is tailored to allow the material to withstand the temperature, pressure, abrasion, and thermal cycling. The chemical composition of a WC-(NiSiB alloy) ceramet powder is a 35 wt.% Ni-Si-B self- fluxing alloy and 65 wt.% macrocrystalline WC with a particle66size range of +46 to -122 µm. The processing parameters are given in Table I. For the FGM linear freeform structures (wall patterns), the WC particle content is varied stepwise from 0 at the bottom to 65 wt.

The processing parameters are given in Table I. For the FGM linear freeform structures (wall patterns), the WC particle content is varied stepwise from 0 at the bottom to 65 wt. % at the top layer of the deposited dense material. For the FGM multiple- dimensional parts, the WC particle content is varied stepwise from 0 at the inside core to 40.6 wt.% at the outside surface of the deposited pars.

P CS M U S F F S y s t e mFig. 1 Schematic diagram of laser-based additive manufacturing system setupTable I Experimental conditions for the rapid manufacturing of FSW-toolsLaser power, W Laser beam diameter, mm Traverse speed, mm/s Powder feed rate of H13 tool steel, g/min Powder feed rate of WC-(NiSiB alloy) ceramet, g/min Offset between two subsequent clad tracks in the horizontal direction, mm Offset between two subsequent layers in the vertical direction, mm Gas flow rate of argon, l/min200-600 1.0 5-15 0.5 –10 1.5-20 0.64 0.38 12-20The standard methods of metallography are applied to obtain microstructural and compositional analyses of the FGMs. All the microhardness tests are carried out using a Vickers microhardness tester and a 500-g load for 15 sec on polished specimens. For microstructural observations, all the samples are selectively etched using a modified reagent consisting of hydrofluoric acid and nitric acid, and then observed using an optical microscope.

For microstructural observations, all the samples are selectively etched using a modified reagent consisting of hydrofluoric acid and nitric acid, and then observed using an optical microscope. A surface texture- measuring device is used to study the surface roughness of the parts. The maximum profile roughness height is the distance between the lowest and highest points on a surface. All the friction-stir welds have been made in a butt-weld configuration using a laser-deposited FSW- tool and an FSW machine as stated in our previous paper [6].67Results and DiscussionRapid Manufacturing of a H13-steel FSW-toolFig. 2 shows the deposited linear freeform structure (wall) of H13 tool steel using a laser power between 250 and 500 watts, at a traverse speed of 8 mm/s to 20 mm/s. The wall is produced by deposition of tracks on top of each other. The feed rate of H13 tool steel powder is 3 g/min. The offset between two subsequent layers on the vertical direction is about 0.38 mm.

The feed rate of H13 tool steel powder is 3 g/min. The offset between two subsequent layers on the vertical direction is about 0.38 mm. A thin layer of the previous clad material is remelted when the subsequent layer is deposited.Fig. 2 Typical features of the deposited layers of the H13-steel wall: (a) surface morphology; (b) cross-section.Fig. 3 shows a laser-deposited H13-steel FSW-tool for the friction stir welding of aluminum alloys. The laser cladding parameters are optimized to be a 350-W laser power, 1.0- mm beam diameter, 10- mm/s traverse speed, 0.64- mm horizontal offset between two subsequent clad tracks, and 0.38-mm vertical offset between two subsequent clad layers. The feed rate of the H13 tool steel powder is 3 g/min. The deposited H13 tool steel exhibits a finer, more uniform martensitic structure with possibly some retained austenite.

The feed rate of the H13 tool steel powder is 3 g/min. The deposited H13 tool steel exhibits a finer, more uniform martensitic structure with possibly some retained austenite. The characteristic rows and layers of the deposited H13 material reflect solidification from the moving laser molten pool as shown in Fig. 3(d). Self- tempering phenomenon at the overlapping regions is observed at some overlapped regions by the extensive heat input as shown in Fig.3(e). The average secondary dendrite arm spacing that is representative of the cooling rate of the deposited H13 metal is about 0.5-1 µm from the microstructural analysis. The calculated cooling rate varies from 1×104 to 1×105 K/s depending on the locations in the deposit of the H13-steel prototypes. The maximum surface roughness of the deposited H13 FSW-tool is about 18 µm. The microhardness of the deposited H13 FSW tool is between 576 and 642 HV0.5.

The maximum surface roughness of the deposited H13 FSW-tool is about 18 µm. The microhardness of the deposited H13 FSW tool is between 576 and 642 HV0.5. The deposited H13-steel FSW-tool exhibits an excellent ability in welding the 6.5 mm thick 6061 aluminum alloy plates as shown in Fig. 3(f). No weld defects are observed in the weld cross-section of 6061-aluminum to itself as shown in Fig. 3(g).68Fig. 3 A laser-deposited H13-steel FSW-tool for the friction-stir welding of aluminum alloys : (a) the as-deposited tool; (b) the tool after hard-turning; (c) cross section of the deposited tool pin; (d) the overlapped laser clad tracks; (e) microstructure at the overlapped region; (f) the tool and the welded 6061 aluminum alloy plate of 6.5 mm in thickness after FSW; (g) cross section of a friction-stir weld of 6061 aluminum to itself.Rapid Manufacturing of a FSW tool of WC-(NiSiB alloy) ceramet/H13 tool steel FGMFig.

4 shows the deposited freeform structures of FGM are deposited using a laser power between 250 and 500 watts at a traverse speed of 8 mm/s to 20 mm/s. The weight fraction of the WC particles varies stepwise from 0 at the bottom to 65 wt.% at the top of the as-deposited wall as shown in Fig.4. The walls are produced by a stepwise deposition of tracks on top of each other. The offset between two subsequent layers on the vertical direction is about 0.38 mm. For the functionally graded walls that start with the pure H13 tool steel powder, the amount of the WC-(NiBSi alloy) ceramet in the powder mixture is increased steadily to generate material with a compositional gradient. The designed distribution of WC particles (wt.%) in the deposited FGM walls is 0, 23.1, 40.6, 54.2, and 65 from the bottom layer to the top layer. The graded wall in Fig. 4a is 40 mm in length and 12 mm in height, consisting of 31 discrete layers. The wall thickness is approximately 1.0 mm.

The graded wall in Fig. 4a is 40 mm in length and 12 mm in height, consisting of 31 discrete layers. The wall thickness is approximately 1.0 mm. At a low magnification, the layered structure of the wall is also visible. The surface contains a number of white WC particles originating from the powder- feeding process. The clad layer consists of the WC particles, ?-Ni primary dendrites, and the eutectics at the interdendritic regions. From the cross section (Fig. 4(b)), the composition of the FGM changes gradually from the pure H13 tool steel at the bottom layer to the pure WC-(NiBSi alloy) ceramet at the top layer. By adjusting the powder variants and the laser processing parameters used, laser cladding process is capable of fabricating crack- free FGM wall patterns that are dense, metallurgically bonded, and geometrically accurate. Fig. 4(h) shows the deposited pattern consisting of several FGM walls. The deposited walls are 20 mm in height and 1.0 mm in thickness.

Fig. 4(h) shows the deposited pattern consisting of several FGM walls. The deposited walls are 20 mm in height and 1.0 mm in thickness. Although the walls are easily made from the pure H13-steel powder, the addition of hard-phase particles like WC brings along a significantly higher cracking susceptibility. Since the unavoidable stresses caused by the thermal cycle of the deposition process accumulate with the clad process, crack- free walls are obtained only when the dimensions of the samples are kept small. Different microstructures are obtained with selected carbide contents as shown in Fig.4(c)- (g). With the increase of the carbide content, a large number of carbide particles remain undissolved or only partially dissolved.

With the increase of the carbide content, a large number of carbide particles remain undissolved or only partially dissolved. As the content of the WC particles increases to 65 wt.%, the WC particles grow significantly coarser as shown in Fig.4(g).The aggregation and bridge connection growth of the und issolved WC particles produce flower- like clusters in the molten pool as shown in Fig. 5. When more WC particles are dissolved or partially dissolved, the chemical composition of the matrix alloy is greatly changed.69From Fig. 5(a)-(c), the matrix alloy consists of ?-(Fe, Ni) primary dendrites and eutectics of ?- (Fe, Ni) plus carbo-borides at the interdendritic regions. The content of the eutectics depends on the dissolution degree of the WC particles. The content of eutectics in Fig.5(c) is clearly lower than that in Fig. 5(a). At the regions with a low content of carbon and boron as shown in Fig. 5(d), no eutectic constituent is observed.Fig.4.

6 shows a laser-deposited FGM FSW-tool for the friction-stir welding of hard metals. One of the major challenges in expanding the application of the FSW process to new materials is the design and rapid manufacturing of powerful tools. A WC-based ceramet/tool steel functionally graded material (FGM) is synthesized to build a FSW-tool in order to weld the hard metals. By controlling the amount of different supplied powders under the optimized laser cladding conditions, the content of WC particles around both the tool pin and the shoulder is varied stepwise from 0 at the inside core to 40.6 wt.% at the outside surface. The hardness of the WC particles, measured at a load of 500 g for 15-s, is between 1845 and 2435 HV0.5 and depends on its partial dissolution, reprecipitation, and growth process. In contrast, the matrix alloy, which is a resolidified products of H13 and the molten WC-(NiBSi alloy), is found to be in the70range of 400 to 540 HV0.5.

In contrast, the matrix alloy, which is a resolidified products of H13 and the molten WC-(NiBSi alloy), is found to be in the70range of 400 to 540 HV0.5. The hardness of the matrix alloy is greatly affected by the content and partial dissolution of the WC particles. The surface roughness of the FGM FSW-tool is between 12.5 and 16.7 µm. The FGM prototypes that have a smaller surface roughness than the H13 tool steel prototypes may be due to the self- fluxing nature of the NiSiB matrix alloy. The feasibility study of using this FGM tool to friction-stir weld the 6.5 mm thick Ti-6Al-4V alloy has recently been carried out as shown in Fig.6(c) and (e). Although some weld defects such as the voids are fo und in the weld, distinct plastic flow of titanium alloys is developed by the stirring action of the FGM tool. This progress is still under development together with the friction stir welding of a 316L stainless steel to itself.Fig.6.

The one-step laser cladding71process is capable of creating a novel, fully dense and metallurgically bonded near- net-shape tool with microstructural and compositional gradients. Two kinds of FSW-tool materials are developed and evaluated for welding different materials. A pure H13 tool-steel powder is used to build a FSW-tool for welding the materials with relatively low melting temperatures, such as magnesium and aluminum alloys. A WC-based cermet/tool steel functionally graded material (FGM) is synthesized to build a FSW-tool in order to weld the materials with higher melting temperatures and highly abrasive materials such as MMC materials. By controlling the amount of different supplied powders under the optimized laser cladding conditions, WC-based cermet/tool steel FGMs are successfully synthesized layer by layer.

By controlling the amount of different supplied powders under the optimized laser cladding conditions, WC-based cermet/tool steel FGMs are successfully synthesized layer by layer. The FGM exhibits an expected gradient structure, a high bonding strength, and a distribution in functionality of synthesized layers between the superhard surface with an excellent wear resistance and the inside core with a combination of toughness and hardness at elevated temperatures.AcknowledgementThis work was financially supported by THECB, Grants 003613-0022-1999, 003613- 0016-2001, NSF Grants No. DMI-9732848 and DMI-9809198, and the U.S. Department of Education, Grant No. P200A80806-98. The contribution of Dr. Dongming Hu is acknowledged.Reference[1] C.D. Sorensen, T.W. Nelson, and S.M. Packer, Tool Material Testing for FSW of High- Temperature Alloys, Proceedings of the 3rd International Symposium on Friction Stir Welding, Port Island, Kobe, Japan, 27-28 Sept. 2001. [2] T.J. Lienert, K.V.

plication of Machine Learning Methods to the Open-Loop Control of a Freeform Fabrication SystemEvan Malone, Oliver Purwin Mechanical & Aerospace Engineering, Cornell University, Ithaca NY USAAbstract Freeform fabrication of complete functional devices requires the fabrication system to achieve well-controlled deposition of many materials with widely varying material properties. In a research setting, material preparation processes are not highly refined, causing batch property variation, and cost and time may prohibit accurate quantification of the relevant material properties, such as viscosity, elasticity, etc. for each batch. Closed-loop control based on the deposited material road is problematic due to the difficulty in non-contact measurement of the road geometry, so a labor-intensive calibration and open-loop control method is typically used.

In the present work, k-Nearest Neighbor and Support Vector Machine (SVM) machine learning algorithms are applied to the problem of generating open-loop control parameters which produce desired deposited material road geometry from a description of a given material and tool configuration comprising a set of qualitative and quantitative attributes. Training data for the algorithms is generated in the course of ordinary use of the SFF system as the results of manual calibration of control parameters. Given the large instance space and the small training data set compiled thus far, the performance is quite promising, although still insufficient to allow complete automation of the calibration process. The SVM-based approach produces tolerable results when tested with materials not in the training data set.

The SVM-based approach produces tolerable results when tested with materials not in the training data set. When control parameters produced by the learning algorithms are used as a starting point for manual calibration, significant operator time savings and material waste reduction may be achieved.IntroductionSolid freeform fabrication (SFF) is the name given to a family of manufacturing processes which allow three dimensional printing of arbitrarily shaped structures, directly from computer-aided design (CAD) data.(a)(b)Figure 1. Fabrication platform: (a) 3-Axis gantry robot for deposition with cartridge/syringe tool, (b) continuous wire-feed toolTypically, an SFF system consists of a tool which dispenses a “road” or line of material, a robotic positioning system which moves the tool along 3-dimensional trajectory, and a software control system. SFF has traditionally focused on producing passive mechanical parts.

SFF has traditionally focused on producing passive mechanical parts. Advances377in this technology and developments in materials science make it feasible to begin the development of a single, compact, robotic SFF system – including a small set of materials – which can produce complete, active, functional electromechanical devices, e.g. mobile robots. A research SFF system with two material dispensing tools has been constructed (Figure 1Error! Reference source not found.) pursuant to this goal, and Figure 2 depicts a Zn-air battery produced with this system. One of the challenges in developing such a system is in achieving precise, accurate, and repeatable dispensing of materials despite the difficulty of automatically monitoring output quality, and the significant variations in properties between materials – even between batches of the same material.

Currently, these challenges are handled via an extensive and laborious manual calibration process for each batch of each material, immediately prior to use. During calibration, the SFF system produces a series of rectilinear test patterns (Figure 7), and control parameters are tuned until the produced pattern matches the desired pattern to the operator’s satisfaction.Figure 2. Zn-air battery produced via SFFThe control parameters (Table I) describe piecewise-linear profiles (Figure 3) for the commanded extrusion rate from the tool and for the robot trajectory speed. The k-Nearest Neighbor and SVM algorithms have been applied to the computation of control parameters for the syringe tool (Figure 1Error! Reference source not found.a). The inputs (i.e. “attributes”) to the learning algorithms are simple quantitative and qualitative descriptions of the material, the tool, and the deposited road (Table II).

The inputs (i.e. “attributes”) to the learning algorithms are simple quantitative and qualitative descriptions of the material, the tool, and the deposited road (Table II). The attributes consist of parameters which are strictly378dependent on the material itself as well as desired properties of the extruded material road (e.g. width and height of the dispensed material). In that these attributes have been intuitively selected, it is unclear whether they are necessary or sufficient to fully represent the problem domain.Table I.

Each successful test pattern (e.g. the material has been dispensed with uniform width over the entire road length, and deposition begins and ends at the start and end, respectively, of the geometric path) is considered a training sample. The data STARTERR and STOPERR, which describe the observed delay between the start of path motion, and the commencement of material deposition, are used to determine acceptable training data. Ideally, only training data for which both STARTERR and STOPERR have the value OK would be used, but given that achieving a single such data point can take more than an hour, STARTERR and STOPERR values of OK, EARLY and LATE are considered acceptable. Although this adds some noise to the data set, it increases the number of data points by an order of magnitude, and it is believed that the noise will be roughly symmetrically distributed about the ideal parameter values.

A training sample contains the material, tool and path attributes and the required control parameters. Using these criteria, the training data set comprises 63 calibrations and 9 separate materials.During the training process the main objective is to generate as many training samples as possible, the actual width and height of the dispensed material is of minor importance, as long as380the road is uniform. On the other hand, when testing the trained machine learning algorithms, the road properties become very important since the objective of the SFF process is to dispense the material with a well-defined geometry.

The precision and accuracy of material dispensing on test patterns, as well as the time required for calibration are used as the basis for comparison between each of the machine learning methods and the manual calibration method.ImplementationTwo different algorithms for the machine learning task were implemented: k-Nearest Neighbor (k-NN) and Support Vector Machines (SVM). The following sections cover the implementation and test of these algorithms. Vectors will be indicated using bold letters. A bold letter with an index denotes a particular vector out of a set of vectors.Description of k-NN ApproachThe k-NN algorithm is a memory-based or instance-based algorithm, meaning that the entire training dataset remains in memory (see Mitchell for in-depth explanation). There is no explicit training step for the algorithm or extraction of a generalized classification rule.

There is no explicit training step for the algorithm or extraction of a generalized classification rule. During classification (generation of a set of control parameters for a new material), the algorithm compares the new material description vector, xnew, with all description vectors in the training set and computes a "distance" function based on the attributes (material, tool, and road description data). The label or parameter vector ynew for xnew is a combination of the parameter vectors of the closest neighbors. The idea is that attribute vectors which are close in the instance space should also have similar labels. This method has the advantage that an entire vector of labels can be mapped to a new example at once as opposed to SVMs, where each SVM can only produce a single output value. Since this algorithm is relatively easy to program, an implementation was written in C as part of the project.

Since this algorithm is relatively easy to program, an implementation was written in C as part of the project. The training and test data are stored in text files to maintain readability by a human user.Procedure to classify a new attribute vector xnew:1) Normalization of the attributes 2) Computation of distances to all samples in the training set 3) Sorting of training set according to distance to xnew 4) Computation of weighted average of the labels of the k nearest neighborsNormalizationIn the problem at hand, different attributes have different ranges. For example, the average "Pushout Speed" is three orders of magnitude larger than the average "Pushout Time". In order to ensure an equal weighting of the distances for all attributes xi, all attribute vectors are normalized to a range of 0 to 1. The normalization is based on the minimum and maximum values xi,min and xi,max of each attribute in the training data.

The normalization is based on the minimum and maximum values xi,min and xi,max of each attribute in the training data. The normalized attribute xi,norm is therefore:381xi,norm=xix−maxx−min , i xi,minThe entire training set and the new examples are normalized.Compute DistancesThe distance d between two attribute vectors x1 and x2 is defined as the Euclideandistanced(xx , 12)=N∑(x,1i−x,2i)2i1 =where N is the number of attributes in the attribute vectors.Sort the training setAfter the distances of the new example to all instances in the training set have been computed, the training set is sorted according to distance.

Since the training set in this particular case is not large (less than 100 examples), a very simple "Insert Sort" algorithm has been implemented.Compute new labelThe label ynew for the new example xnew is calculated according toynew=k∑ ∑yw ii1 = kw ii1 =iwhere k is the number of neighbors to be considered, yi are the labels of the closest neighbors, and wi are the weights assigned to each neighbor. There are many different ways of defining the weights wi, but in general closer neighbors should carry more weight than neighbors who are farther away. Two different weight functions were used. The first attempt was based on the inverted distance to the new example:w i=d(x1new,xi)The distance d had to be checked before computing the weight in order to avoid a division by zero in case xnew matches up perfectly with one of the instances from the training set.

This seemed to work fairly well, although large numbers of neighbors increased the leave-one-out error.The second version used an exponential drop-off:w i=exp[d(x1new,xi/)f]382where f is the kernel width, a parameter to control the amount of roll-off. The second version leaves more room for tuning, since the parameter f can be chosen freely.ExperienceThe k-Nearest Neighbor algorithm is very easy to use, and it is straightforward to understand exactly how a particular classification was made. The algorithm is very fast given the small data set. Furthermore, there are no convergence problems (see also the SVM section).In general the quality of the output data is good but not great. There are no cases of blatant misclassification, since it looks for labels that have been used before and combines them. On the other hand, that is one of its biggest drawbacks: k-NN in the current version cannot extrapolate beyond what it saw before.

On the other hand, that is one of its biggest drawbacks: k-NN in the current version cannot extrapolate beyond what it saw before. If a new material or tool configuration is used that's outside the “known region” of the instance space of the training data, the results will be poor.One particular problem with the data is that it comes in clusters. During each manual calibration run, five or six training samples were collected. These attribute vectors/labels don’t differ much from each other, since they have the same material and the same nozzle. The k-NN tends to lock onto the closest cluster in the instance space and produce labels that are very similar to the ones in that particular cluster, instead of averaging over several clusters. This was especially apparent when using the inverse distance as a weighting function. Small numbers of neighbors produced lower leave-one-out errors, but didn't do much averaging, i.e. "learning".

This was especially apparent when using the inverse distance as a weighting function. Small numbers of neighbors produced lower leave-one-out errors, but didn't do much averaging, i.e. "learning". In that case, the algorithm basically works as a database which looks up the closest instance in the training set. On the other hand, large numbers of neighbors generalized too much and the differences between labels started to blur. All classifications looked more or less the same, which leads to very mediocre results.The exponential weighting method is superior, especially when using a large number of neighbors.

All classifications looked more or less the same, which leads to very mediocre results.The exponential weighting method is superior, especially when using a large number of neighbors. Better results were achieved by setting k to the total number of examples in the training set and just tuning the kernel width f.Description of SVM ApproachSupport Vector Machines operate on the principle of calculating an optimal separating surface between data points with different labels, such that similarly labeled data all (or nearly all) lie on one side of the separating surface. SVMs can be used not only for identifying discrete classes, but also can be applied to regression problems, i.e. finding the function f: Xf→ N ,Xℜ∈YYℜ∈given the training data setDxi=∈{ ( 1 , yXx,iy 1 ∈),K Y(,xl,yl} )383where N is the dimension of the instance space and l is the number of training examples.

A good introduction to this topic can be found in Smola and Schölkopf. The regression takes the formf(xnew)y ==xw •new+bThe vector w is the weight vector, b is the offset. In order to find w and b the following optimization problem has to be solved:min1 2w2+Cl∑i1 =( ξξ + ii)subject to −yixw •i+≤−bξε ixw •ib −+yi+≤ξε i, ξξ i i≥0being the slack variables. The with C being the weight on the slack variables and parameter ε represents the distance from the linear classifier (as defined by w and b) up to which values of y , then additional cost is added to the minimized function. This also explains why there are two slack variables: one is for the positive, the other one for the negative deviation from the linear classifier. In practice, the dual of this problem is solved and the regression takes the form, i ξξ ii are considered "good".

In practice, the dual of this problem is solved and the regression takes the form, i ξξ ii are considered "good". If yi differs by more than ε fromb+• xwy=l∑xxαα ( ) i newxxαα ( ) i i)+i1 =where l is the number of support vectors, αi and αi* are the Lagrange multipliers of the trained SVM, xi are the support vectors, and b is the offset. The function k(.) is the kernel.ImplementationAn existing C-library SVM implementation (SVMlight, see Joachims) is used for the training and regression. A wrapper function was written that handles the input and output to SVMlight, i.e. it prepared the input files and read the data from the output files, as well as making function calls to the SVM library with the appropriate parameters, e.g. the kernel type.As opposed to kNN, SVMs in regression mode have only one output. In order to compute the full control vector y with 12 components, 12 SVMs were run in parallel.

the kernel type.As opposed to kNN, SVMs in regression mode have only one output. In order to compute the full control vector y with 12 components, 12 SVMs were run in parallel. This approach is justified, since the entire control law will be modeled by the SVM as a black box. Therefore any coupling between the control parameters will implicitly be taken care of. The full regression is thus:384y=⎡ ⎢ ⎢ ⎢ ⎣y 1M y 12⎤ ⎥ ⎥ ⎥ ⎦=⎡ ⎢ ⎢ ⎢ ⎣xw • 1w12xnewMnew++b 1b 12⎤ ⎥ ⎥ ⎥ ⎦where wi and bi are the weight vector and offset of the ith SVM. All SVMs were trained on the same attribute vectors xi, but different labels yi (depending on the control parameter). Note that the labels yi are scalars.For some reason, the built-in “leave-one-out error” (see below) computation in SVMlight did not work for regression and so the wrapper function includes that functionality.

The wrapper function reads the entire data set, extracts the line with the example that is to be left out and writes both the single example and the rest of the training data to two different files. Then the SVMs are trained, the example classified and the output compared with the original label.ExperienceThe experience with SVMs was very good. The trained SVMs are able to produce control parameters with small leave-one-out errors and tolerable results when testing them on the physical machine. SVMs are remarkably superior to k-NN when dealing with novel materials and/or nozzles.One observed problem was the convergence of the SVMs when training them for a leave- one-out error computation. It turned out that SVM light is very sensitive to different values of C or higher order Kernels. Usually the training of a single SVM takes only a fraction of a second, but from time to time the algorithm just failed to converge and kept running until the user stopped it.

Usually the training of a single SVM takes only a fraction of a second, but from time to time the algorithm just failed to converge and kept running until the user stopped it. Adjusting the parameter "-e" (error for termination criterion) helped in most cases.ResultsA standard method to compare the performance of machine learning algorithms is to compute the leave-one-out errors. In this test, one data sample gets removed from the training data and the machine learning algorithm is trained on the remaining data. The trained algorithm is then used to classify the left out example and the results are compared to the original label. Ideally, the difference should be zero. This test is repeated for each sample in the data set. The average error is the leave-one-out error.

Dev. 0:23 0:43 4.94Figure 7. Test patterns produced by automatically generated control parameters387ConclusionThe time and effort required to gather data for these experiments, hence the small training data set, has limited the statistical significance of any conclusions. Qualitatively, the SVM algorithm is superior to k-NN with materials not in the training data set. This qualitative observation is supported by the leave-one-out errors, as seen in Figure 4, Figure 5, and Figure 6. Experience of the operator with manual calibrations showed that parameters 2, 6, 7, and 10 are the least important. These are the ones where k-NN has the lowest errors, whereas the SVM has smaller errors for those parameters which have larger impact on the manual calibrations.

These are the ones where k-NN has the lowest errors, whereas the SVM has smaller errors for those parameters which have larger impact on the manual calibrations. As Figure 7 indicates, the k-Nearest neighbor algorithm often generated control parameters that failed to generate any discernable material deposition at all, while the SVM always produced results at least as good as the first guess of a human operator. For the time being, using learning algorithms to produce control parameters has brought no benefit to the deposition performance of the system, and does not yet allow for complete automation of the calibration process. On the other hand, employing the learning algorithms to produce an initial guess at control parameters seems to decrease the time required to obtain satisfactory manual calibration by as much as a factor of 2 (Table III). If this bears out in more extensive testing, the savings in operator time, not to mention chemical waste are very significant.

If this bears out in more extensive testing, the savings in operator time, not to mention chemical waste are very significant. This benefit alone warrants permanent inclusion of the learning algorithms into the software of the system. It remains to be seen whether or not, given a sufficiently large training data set, the learning algorithms will be able to produce completely satisfactory or superior deposition performance without manual intervention.Future work will include the exploration of intelligent weighting schemes for attributes, such as described by Domingos. These will be explored in the context of ongoing use of the system. Such automated weighting schemes have the added benefit of indicating which attributes are most significant in generating classifications, guiding the selection of additional attributes, and removal of redundant attributes. A neural network will also be explored as an alternative to the learning algorithms explored here.References Anguita, D., A.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAPPLICATION OF SCHLIEREN TECHNIQUE IN ADDITIVE MANUFACTURING: A REVIEWR Bharadwaja*, Aravind Murugan*, Yitao Chen*, Dr. F W Liou*Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 654011. AbstractAdditive manufacturing has gained a lot of attention in the past few decades due to its significant advantages in terms of design freedom, lower lead time, and ability to produce complex shapes. One of the pivotal factors affecting the process stability and hence the part quality is the shielding gas flow in additive manufacturing. As extremely beneficial for the process, the shielding gas flow is often set at maximum supply to achieve enough gas cover over the substrate. This causes excessive quantity of shielding gas to be unutilized.

This causes excessive quantity of shielding gas to be unutilized. Realizing the importance of shielding gas, various studies have been carried out to monitor and visualize the shielding gas, and one such technique is Schlieren imaging. Schlieren visualization has been used since the 1800s as a powerful visualization tool to visualize fluctuations in optical density. The Schlieren technique is highly effective for visualizing and optimizing shielding gas flow. This paper aims to provide an overview of Schlieren technique used for visualization of shielding gas and highlights the application of Schlieren in additive manufacturing.2. IntroductionAdditive Manufacturing (AM) is the process of joining materials to make objects from 3D model data, usually layer upon layer [1] and has been in use since the 1980s. AM technology has since undergone over three decades of growth and is currently one of the world's fast-growing advanced manufacturing techniques [2].

AM technology has since undergone over three decades of growth and is currently one of the world's fast-growing advanced manufacturing techniques [2]. The development in additive manufacturing over the past three decades can be classified into three major phases. In its initial phase, it was used as a tool to create prototypes and models, mainly utilized by product designers. AM's second developmental stage involves its application to create completed components (also termed as ‘direct digital manufacturing’). In the third stage, 3D printers are used widely available for use to direct customers at a reasonable price point [3]. As material is deposited layer upon layer, additive manufacturing can produce components of complex parts which cannot be easily manufactured by using conventional methods of manufacturing [4].Schlieren technique has been in use for about two centuries for imaging and analyzing phenomena in transparent media.

The Schlieren method has been commonly used to visualize gas flows in varied applications such as vehicle aerodynamics, ballistics, ventilation methods etc. [5]. This method is particularly helpful for imaging as it generates a natural, easy-to-interpret image of refractive index-gradient fields. Shadowgraph and interferometry are the companions of Schlieren technique, but Schlieren has proven to be more sensitive and better suited for quantitative visualization [6]. Over the years, the Schlieren technique has undergone significant changes, and the latest technology has enabled Schlieren to be applied in advanced research topics. With1537advanced softwares and digital imaging techniques, the capability of the Schlieren technique has been significantly strengthened. Schlieren has recently been used to visualize the shielding gas flow in various advanced manufacturing techniques.

Schlieren has recently been used to visualize the shielding gas flow in various advanced manufacturing techniques. Several vital findings related to nozzle design, gas flow and part quality have been made through the application of Schlieren technique. This paper provides an overview of the Schlieren technique, it’s evolution and application of Schlieren technique in AM.3. Discussion3.1 Evolution of SchlierenSchlieren optical imaging system was introduced in the 1800s by Robert Hooke [5]. Schlieren optics proved to be a sound technique for visualizing transparent media [7]. With his rudimentary instruments, Hooke explained the light refraction due to density gradient in his book Micrographia [8]. The knife edge, which is now an important part of the setup to cut off the diffracted rays and increase the sensitivity of the device was introduced by Leon Foucault in the mid-19th century [5,9].

The knife edge, which is now an important part of the setup to cut off the diffracted rays and increase the sensitivity of the device was introduced by Leon Foucault in the mid-19th century [5,9]. August Toepler, between 1859 and 1864, reinvented Schlieren technique and was given official recognition for this. He studied and proposed theories of Schlieren technique and provided in-depth procedures for designing a Schlieren system [10].3.2 Basic concepts of SchlierenThe Schlieren technique is based upon the Snell's Law [5,10], which states that light travels with different speeds in different media. While encountering different media in its path, light rays get refracted, and deflect from their original path resulting in Schlieren images [10]. The refractivity (n-1) of any fluid is given by:n-1=κρ……………………………Equation (1)where, n is the ratio of the speed of light in vacuum (3x108 m/s) to the speed of light in the medium.

The refractivity (n-1) of any fluid is given by:n-1=κρ……………………………Equation (1)where, n is the ratio of the speed of light in vacuum (3x108 m/s) to the speed of light in the medium. κ is Gladstone Dale coefficient and is about 0.23 cm3/kg of air in standard conditions and ρ is the density of the fluid.In a Schlieren setup, (considering a Z type setup as an example), the light produced from the light source passes through one mirror, enters the given medium and then gets refracted. This light beam again passes through the second mirror and reaches the knife edge. As the image has been diffracted by the medium, there would be a shift from its original position as shown in Figure 1. The expression for the shift is given by:∆a = f . ε…………………………..Equation (2)where, f is the focal length of the second mirror and ε is linear in the y component of refraction angle E in the Schlieren object [5].

The brightness of the picture obtained from the Schlieren technique is determined by the slit orientation as shown in the Figure 2. There are three orientations of the slit available: vertical, horizontal, and iris. According to Siewert et al. [12], the iris can be used to visualize density gradient in all directions, but the images are characterized by a lower brightness of the image. Evaluation of the geometry influence and slit orientation makes it clear that images with high clarity can be obtained with slits perpendicular to the work piece. It is also recommended that, when working with identical concave mirrors, the shape of the light source must match with the shape of the slit opening. While using an iris as a slit, an LED as a light source would give best results (as LED produces circular beam of parallel light) and while using a vertical slit opening, an elongated rectangular light source is preferable.Figure 2: Impact of slit orientation on image brightness.

Reproduced with permission from Welding Journal, (January 2014), © American Welding Society15393.3.2 Slit SizeIn a Schlieren setup the amount of light entering the system is governed by a slit placed just in front of the light source. With the reduction of the slit width, a lower amount of light would enter the system and hence lesser amount of light would be diffracted. This leads to visualization with lesser variation between density of the components. The sensitivity and measurement range of the Schlieren system are determined by the size and shape of the source slit [5]. The main objective of the slit size variation (specifically in the applications of Schlieren in additive manufacturing processes) is to visualize the type of flow of the shield gas from the nozzle [12]. Siewert et al observed that with the decrease in slit size, the radiation from the arc decreases as depicted in Figure 3 [12].

Siewert et al observed that with the decrease in slit size, the radiation from the arc decreases as depicted in Figure 3 [12]. It can be noted that the brightness of the welding arc increases as the slit size increases (from left to right), because with increase in the slit size, greater amount of light enters the system.Figure 3: Schlieren aperture slit of 2 × 6 mm (left), 3 × 6 mm (middle), and 5 × 6 mm (right). Reproduced with permission from Welding Journal, (January 2014), © American Welding Society3.3.3 Color SchlierenColor Schlieren technique has been widely used since 1985, and it plays a vital role in visualization of the shielding gas in the recent times [13]. Significant advancements in technology has made color Schlieren more robust. Wolter et al. notes that color Schlieren images signifies vector displacements and the angle in the cutoff plane whereas grayscale images are limited to signifying the scalar quantity [14].

Wolter et al. notes that color Schlieren images signifies vector displacements and the angle in the cutoff plane whereas grayscale images are limited to signifying the scalar quantity [14]. The color has an added advantage of sensitivity and hue saturations and color intensity, that can be used for quantitative analysis of the image. In grayscale images, the contrast is uneven while the color images are evenly balanced. It offers additional dimensions when needed to encode the characteristics and enhance the contrast of the data presented. Color Schlieren uses a color-grid cutoff in the place of the traditional knife-edge to produce images with hue variations in place of grayscale values [15]. According to Siewert et al [12], among the four filters of colors (blue, yellow, red, and green), the pair of two colors: blue/yellow and red/green makes the best pair.

According to Siewert et al [12], among the four filters of colors (blue, yellow, red, and green), the pair of two colors: blue/yellow and red/green makes the best pair. These enabled turbulences to be visualized with a greater contrast as shown in Figure 4.1540Figure 4: Schlieren images, used filter pairs: blue/yellow (left) and red/green (right). Reproduced with permission from Welding Journal, (January 2014), © American Welding Society3.4 Recent Developments in Schlieren Technique3.4.1 The CameraTraditional cameras with high speed drum, bulky size, rotating mirror and rotating prism have been replaced by the modern digital high speed cameras, reaching up to 2000 frames per second (fps) and with CMOS sensor capabilities [16]. Schlieren technique that’s completely based on visualization has also undergone a technological shift.

Schlieren technique that’s completely based on visualization has also undergone a technological shift. The technique which was resolute to visualization of simple candle plumes is now able to capture images of shock waves produced by a bullet and sonic booms produced by a supersonic airplane. The new era of Digital Single Lens Reflex (DSLR) have completely replaced the SLR cameras with their ease of use and their image clarity. The light source being highly intense in a Schlieren setup, Schlieren images do not need the same exposure that a traditional photograph image would require [16]. Also, the SLRs required a film to store the image and a dark room to publish a photograph. To eliminate this time loss, DSLR comes up with immediate results of the test shots making the Schlieren setup more user friendly. The color balance that once needed numerous films, can be now performed right through the camera.

The color balance that once needed numerous films, can be now performed right through the camera. A wide variety of interchangeable lenses are available for the DSLR cameras making the camera more efficient and the lenses have outboard focusing feature which eliminates the use of additional lens.3.4.2 LED as a light sourceThe light source for the Schlieren technique consisted of spark gaps, lasers, xenon tubes, arc lamps and incandescent filament lamps in the 20th century [5]. Light Emitting Diode (LED) was in its nascent stages of development back then. By 2000s, LEDs had advanced tremendously and was being used for numerous scientific applications [17]. The LEDs stood as a non-coherent light emitter and an inexpensive light source alternative for Schlieren technique. The narrower the light source is, the sharper would be the image produced [16], and hence LED stands as a great option, as it can effectively produce a light beam of 15o.

The narrower the light source is, the sharper would be the image produced [16], and hence LED stands as a great option, as it can effectively produce a light beam of 15o. Focusing it directly onto the first Schlieren lens or the mirrors can eliminate the need for condenser and a slit [18]. Spark gaps, which were considered as a light source earlier were hazardous to use, more expensive and not durable [19]. The LED source offers both alignment and high-speed imaging pulsed output. Only a laser source could provide a pulse width enough to freeze the motion earlier for PIV and Schlieren velocimetry1541applications [16]. However, lasers posed eye safety hazards and hence LEDs replaced them as a reliable and robust light source for Schlieren imaging.3.4.3 Background oriented SchlierenThe most significant recent developments in this domain is certainly the background- oriented Schlieren.

Unlike a conventional Schlieren setup, background oriented Schlieren eliminates the use of expensive mirrors, lasers and the knife edge [20]. It uses a background with a lot of fine detail. Light from the light source falls on the background film, which then passes through Schlieren S ( any substance which has a different density than air and is able to diffract the light ) placed in the test area as shown in Figure 5 and is then visualized using a camera. Using digital imaging software, the two images (one with Schlieren S placed and the other without) are compared and the disturbances caused by the diffraction are captured [16]. A small spot on the background seen as the dashed line in Figure 5, is shifted by refraction by the media. The shift ∆y can be calculated by comparing images of the background, with and without the medium in the test area.

The shift ∆y can be calculated by comparing images of the background, with and without the medium in the test area. The deflection angle can be calculated after measuring ∆y and hence the density of the medium is known.Figure 5: Schematic diagram of Background Oriented SchlierenAs most portions of the traditional Schlieren setup are eliminated, the background oriented Schlieren is inexpensive, small, and accurate. This technique can be combined with particle image velocimetry (PIV) as well to produce visualizations [21]. Another added advantage that this technique has is that it produces a field of view which cannot be achieved in conventional Schlieren methods.

Another added advantage that this technique has is that it produces a field of view which cannot be achieved in conventional Schlieren methods. The field of view can be at least half the size of the background [16] and large objects like helicopters or supersonic aircraft can be visualized using wide natural background [22] as shown in Figure 6.1542Figure 6: Use of Schlieren technique for measuring the strength and position of shockwaves caused by supersonic planes. Image reproduced from NASA.3.5 Schlieren In Additive ManufacturingNon-thermal atmospheric-pressure plasma jets are gaining a lot of attention as a promising tool for additive manufacturing [23]. Instead of using a large nozzle for the shielding gas, micro capillaries have emerged as alternatives for localized processes. The dynamics of gas outside the capillaries and the behavior of the charged particles towards the inert gas are pivotal in deciding the efficiency of the capillaries.

The dynamics of gas outside the capillaries and the behavior of the charged particles towards the inert gas are pivotal in deciding the efficiency of the capillaries. Hence, to design and control such micro plasma jets, a closer look on the functioning and understanding the plasma and gas-surface interaction is essential.As the micro jets are invisible to the human eye or a normal DSLR, Schlieren technique is being used to determine the flow from the micro plasma jets [23]. In the study by Bradley et al. [23], it was observed from the Schlieren images that the micro jets maintained a laminar flow up to 54 mm from the nozzle and then the outflow turned into turbulent. The regions of laminar and turbulence flow can be easily distinguished as indicated in Figure 7. Hence Schlieren technique can be effectively used in applications which require the type of flow to be studied.Figure 7: Schlieren image showing the variation in flow from a nozzle.1543Anilli et al.

in [24] used selective laser melting (SLM) to produce single and double chambered laser cutting nozzles. SLM nozzles can be lightweight, have inner channels and can be created in any complex shapes. Nonsymmetrical nozzles have also gained attention in recent years and these can’t be manufactured by using conventional manufacturing methods and hence additive manufacturing stands as a solution to this problem [25]. Anilli et al. used gas atomized 18Ni300 maraging steel powder for the producing nozzles using additive manufacturing for the experimentation. A conical convergent single chamber and corresponding double chamber nozzles of 1 mm size were manufactured as shown in Figure 9. The gas flow through these nozzles was then tested using Schlieren imaging technique.

The gas flow through these nozzles was then tested using Schlieren imaging technique. The setup consisted of five mirrors, diffracting more light than a conventional z type Schlieren setup as shown in Figure 8.L E O S o u r c eM i r r o r 1M i r r o r 2M i r r o r 3M i r r o r 4M i r r o r sFigure 8: Schematic diagram of the Schlieren setup.Figure 9 Axial section of conical convergent single chamber (a) and corresponding double chamber nozzle (b)1544The light source which was chosen for the experimentation was a white LED. With the help of Schlieren imaging, the gas flow dynamics in stationary conditions were studied at 15 bar of pressure. Three nozzles: standard, SLM as built, and SLM post processed were tested using Schlieren technique.It was observed in a standard nozzle that a bell-shaped shock wave front which collapses to a transverse shock wave, making the first part of the flow separated.

Considering the SLM as built nozzle, the bell shape was maintained and had retracted closer to the nozzle. This was due to the irregularities in the inner surface of the nozzle and irregular profile of the manufactured hole. Hence to achieve a smoother profile, the nozzles had undergone finishing through electrochemical machining (ECM) and abrasive flow machining (AFM). In the case of the post processed nozzle, whose inner surfaces were smoothed also with a more regular hole size, the flow resembled that of the standard nozzle. From the results of Schlieren visualization, it was concluded that nozzles made from SLM required post processing. The inner surface and the roundness of the hole play a substantial role in nozzle production. Here Schlieren technique played a crucial role in determining the optimal nozzle specification, the nozzles that had been subjected to post processing have met the desired performance.

Here Schlieren technique played a crucial role in determining the optimal nozzle specification, the nozzles that had been subjected to post processing have met the desired performance. It was also found that AFM using SiC as abrasive particles was an adequate solution for improving nozzle tip quality as it reaches the effective processing pressure at the nozzle tipShielding gases are fundamental components of several arc welding components. Though the main role of the shielding gas is to protect the melt pool against atmospheric gases, it also acts as a medium for the flow of electrical current between the electrode and the workpiece in the plasma jet. Though most of the shielding gases consist of inert gases such as argon, adding other gases such as carbon dioxide, oxygen or helium, improves arc formation, heat transfer, material weld ability or gas coverage [26].

Several other benefits such as refinement of weld structure and thus improving the mechanical properties of the weld were noted by pulsing the shield gas [27,28]. The dynamics of melt pool are affected by alternating the shield gases [29]. The effects of alternating argon and helium in Gas Tungsten Arc Welding (GTAW) were visualized through Schlieren imaging, and the arc characteristics and flow rates were analyzed by contrasting the image sequences of pure and premixed gas flows [26].Bitharas et al. in [26] studied the use of alternating shielding gas flow in GTAW. In this study, it was noted that pure argon diffracted lesser light than helium. Being heavier than air, argon at 5 l/min spread out from the nozzle and covered the work piece, hence preventing it from oxidation.

Being heavier than air, argon at 5 l/min spread out from the nozzle and covered the work piece, hence preventing it from oxidation. At a higher flow rate of 10 l/min, argon formed eddies near the workpiece mixing with the surrounding air.In the case of pure helium, large dark patches were found in Schlieren images in hot as well as cold conditions. Unlike argon, lesser vortex formations were observed in the case of helium due to higher viscosity and greater buoyant forces. Helium comparatively covered larger surfaces of the workpiece than argon. 50% duty cycle which refers to switching argon and helium alternating for equal intervals of time was used at the time of experimentation. Two switching sequences: 2 Hz and 8 Hz were investigated in this study.

Two switching sequences: 2 Hz and 8 Hz were investigated in this study. With the help of Schlieren technique, it was observed that, images of 8 Hz were brighter than images of 2 Hz.1545As it can be noted from the above examples: the type of gas flow, nozzle design and shielding gas, are all important parameters for additive manufacturing applications. Schlieren technique hence can be used as an effective means to monitor and control these parameters to produce a higher quality additive manufactured component.4. Summary and ConclusionThis paper provides an overview of the Schlieren technique, evolution of Schlieren, and application of Schlieren in AM. Though Schlieren has been in practice since the 1800s, the latest digital technology has propelled the rapid evolution of this technique. Orientation and slit size are amongst the pivotal factors in Schlieren that influence the clarity of the image produced.

Orientation and slit size are amongst the pivotal factors in Schlieren that influence the clarity of the image produced. With the advent of DSLR cameras, the experience of Schlieren imaging has turned more robust and user friendly. LEDs are being considered as a viable alternative to the conventional arc lamps and laser sources due to their ability to act as a narrow light source, hence producing images with better clarity. Schlieren imaging has been utilized to determine the type of flow, nozzle design applications and alternating shielding gases - all of which are important parameters from additive manufacturing standpoint. Hence, Schlieren can be used as a powerful tool for controlling key parameters to produce high quality components for additive manufacturing applications.References1. J. Alcisto, A. Enriquez, H. Garcia, S. Hinkson,T. Steelman, E. Silverman, P. Valdovino,H. Gigerenzer, J. Foyos, J. Ogren, J. Dorey, K. Karg, T. McDonald, and O.S.

PLICATION OF SFF TO PREOPERATIVE PLANNING AND SURGICAL REHEARSAL FOR TREATMENT OF LIMB DEFORMITIES IN DOGS.Ola Harrysson1, Denis Cormier1, Denis Marcellin-Little2, Ketan Jajal11 Department Of Industrial Engineering, North Carolina State University, Raleigh NC 27695,2 Department of Clinical Sciences, North Carolina State University, Raleigh NC 27606AbstractThis report describes SFF-aided assessment and preoperative planning for treatment of bilateral multifocal pelvic limb deformities in a one-year old German Shepherd dog. CT scans were converted to solid models and stereolithography patterns were prototyped. RTV silicone molds were constructed and three sets of polyurethane patterns were cast for pre-surgical rehearsal. The paper compares traditional osteotomy planning procedures, planning via virtual prototypes, and planning with full-scale physical prototypes.IntroductionBone deformities are relatively common in both humans and animals.

Most deformities are developmental and are treated early, in skeletally immature patients. Adults also have deformities that may result from metabolic diseases, malunions, or untreated developmental deformities. Corrective osteotomies are generally performed in one or more locations to treat these deformities. A precise and accurate preoperative planning is critical to the success of these procedures and is difficult to achieve using radiography and 2-dimensional computed tomography (CT).A one-year-old female German Shepherd dog was referred from Charleston, SC to the orthopedic service of the College of Veterinary Medicine at North Carolina State University (NCSU) in November 2001 for evaluation and treatment of multifocal deformities of the pelvic limbs. The dog had angular and rotational deformities affecting femora and tibiae.

The dog had angular and rotational deformities affecting femora and tibiae. Due to the unusual nature of these deformities and their geometric complexity, the information collected from the orthopedic examination and from radiographs of the pelvic limbs was considered to be insufficient for preoperative planning. In February 2002, the orthopedic clinician (DJM) learned about the medical modeling capability available in the Industrial Engineering Department at NCSU. Biomodeling was first fully developed during a European project called PHIDIAS [1-3]. Biomodels have been used for over a decade to diagnose, plan and rehearse medical procedures on human patients with great success [4-7]. Biomodels have been used in the veterinary medicine to aid in reconstruction of skull deformities but, to our knowledge, biomodels have not been used to assess deformed limbs and plan and rehearse corrective osteotomies.

A CT scan of the pelvis and pelvic limbs was performed upon return of the dog to NCSU, and three full-scale biomodels were prepared.17Figure 1 - Rear View of German Shepherd Showing Bone DeformitiesMaterials and MethodsA CT-scan was acquired at the radiology department using a General Electric Computed Tomography scanner. The scan was performed at 0 gantry tilt, and helical scanning was used for faster and more accurate images. The helical scan was reconstructed into 1 mm slices with a resolution 512x512 and a total of 566 images were produced covering the dog from pelvis to extremities. Because of the unusual scanning protocol and the large size of the region scanned, three trials and GE technical support were needed to acquire optimal images. The reconstructed CT-data was transferred to a CD and loaded into the Mimics software (Materialise, Belgium) [8] that is used at the Rapid Prototyping Laboratory at the Industrial Engineering Department at NCSU.

The final scan was performed in two separate sections with a waiting time in between. Unfortunately, the patient moved slightly during the wait, thus causing a misalignment of the lower limbs in the CT data. In Mimics, the images were edited one by one to provide a solid model without interior cavities. Through region growing, a 3 dimensional model of the dog’s hindquarters was made.18Figure 2 - 3D-model dog's pelvic limbs created using MimicsTo fabricate the model, an SLA-190 from 3-D Systems was used with a build envelope of 190 x 190 x 250 mm. Due to the shift in the tibial portion of the computer model, the computer model was separated into two sections that were exported to the STL file format. To fit the model onto the relatively small build platform, the STL files were imported into Geomagic Studios by Raindrop Geomagic [9] and divided into a total of 11 sections.

To fit the model onto the relatively small build platform, the STL files were imported into Geomagic Studios by Raindrop Geomagic [9] and divided into a total of 11 sections. Three build platforms were prepared using the Lightyear software from 3D Systems, and the build style was set to QuickCast to save time, resin and laser hours. Three sets of solid models were requested by the surgeon to plan and rehearse the surgical procedure. One model was to be used to plan the complicated surgery. The second model was to be kept for reference in the operating room. The third model was created just in case an extra model was needed during the pre-surgical planning session. In order to produce the three models, one-piece RTV silicon rubber molds were produced using the QuickCast patterns. A total of seven silicon rubber molds were needed to produce the different bones.

In order to produce the three models, one-piece RTV silicon rubber molds were produced using the QuickCast patterns. A total of seven silicon rubber molds were needed to produce the different bones. The three sets of bones were produced, assembled, finished and painted to give a natural look (Figure 3).19Figure 3 - Silicon molds and polyurethane parts of dog modelThe right pelvic limb was operated on first. Based on the radiographs and the images from the CT-scan, the surgeon had anticipated performing one femoral and two tibial osteotomies to correct the deformities of the right pelvic limb. When the models were presented to him, it was concluded that only the tibial osteotomies would be needed. The pre-surgical rehearsal was performed the day before surgery. The original plan was to perform and opening wedge osteotomy and to use an hinged circular external fixator to correct the proximal end of the tibia.

The original plan was to perform and opening wedge osteotomy and to use an hinged circular external fixator to correct the proximal end of the tibia. A closing wedge osteotomy with a fixating plate was to be used to correct the distal end of the tibia. After working with the model for some time while designing and assembling the external ring fixator, the surgeon decided to correct the deformity by only performing one osteotomy using the external ring fixator at the proximal end of the tibia. Using the model, it was concluded that a medial angular correction of 22 degrees was needed along with a medial rotational correction of 11 degrees.

Using the model, it was concluded that a medial angular correction of 22 degrees was needed along with a medial rotational correction of 11 degrees. Following assembly of the ring fixator, the entire surgical procedure was performed using one of the models, and notes were taken to assist the actual surgery the following day.Figure 4 - Pre-surgical rehearsal on biomodelThe first step of the pre-surgical procedure was to attach the frame’s three rings to the tibia using wires that were drilled through the bone and then attached to the frame with wire fixation20bolts. The wires were later tensioned to increase their stiffness. Half-pins were added to increase the stability of the fixator. The half-pins were threaded into the bone and bolted on the rings. After the external fixator was securely attached to the tibia, the bone was cut between the two upper rings. The distal portion of the tibia was rotated medially 11 degrees.

After the external fixator was securely attached to the tibia, the bone was cut between the two upper rings. The distal portion of the tibia was rotated medially 11 degrees. The frame was then adjusted angularly 22 degrees to check if desired correction had been achieved. The objective of the two corrective adjustments (22 degree wedge and 11 degree rotation) was to achieve proper alignment between the dog’s hip, knee, and foot. After concluding that the practiced procedure did, in fact, produce the desired result, a second frame was assembled and prepared for surgery the following day.The next day, the German Shepherd was prepared for surgery, and the surgical team was briefed about the planned procedure using the biomodel. Two models were used in surgery as references - the one with the frame and one that was untouched.

Two models were used in surgery as references - the one with the frame and one that was untouched. The surgeon performed the surgery following the exact procedure rehearsed the previous day and frequently consulted the models for additional information.Figure 5 – Biomodels in the OR (L) and Attached Ring Fixator (R)Despite the fact that a 22-degree wedge was ultimately needed to achieve the desired result, it was not possible to produce this extreme correction entirely during the surgery. This is due to the fact that muscles, ligaments, and tendons can only be stretched so far before they will tear. The external ring fixator is hinged on one side and has an adjustment screw on the other side that allows the surgeon to gradually increase the wedge angle over a period of time (i.e. a progressive osteotomy).

a progressive osteotomy). In this case, an adjustment of 1-2 degrees per day was made until the final 22-degree angle was achieved.After the final adjustment was completed, the dog was sent home with the external fixator in place to allow the bone to heal back together. After two months, the dog returned to the Veterinary School to have the frame removed. The entire procedure was then repeated to straighten the dog’s left rear leg. A new pre-surgical rehearsal was conducted using the second biomodel. The deformity of the left tibia was very similar to the right, and the same procedure21was planned and practiced. The most significant difference in the two surgeries was that the left femur (thigh bone) had a deformity that required a closing wedge osteotomy with a fixating plate. After the rehearsal, it was concluded that the left tibia also needed a 22-degree angular correction but that no rotational correction was necessary.

After the rehearsal, it was concluded that the left tibia also needed a 22-degree angular correction but that no rotational correction was necessary. The actual surgery was performed the following day, and the biomodels were once again used as references during surgery. Measurements were taken directly on the model to determine where to make the incisions and where to locate the osteotomy sites. The surgery followed the exact same procedure developed during the rehearsal the previous day. The progressive angular correction is currently taking place, and the German Sheppard will wear the external fixator until the end of August 2002.ResultsThe overall outcome of the project was a success, and many lessons were learned. The decision to used QuickCast models to create the silicon molds turned out to be very helpful, since the SLA-190 is a very slow machine that does not have a recoating blade.

The decision to used QuickCast models to create the silicon molds turned out to be very helpful, since the SLA-190 is a very slow machine that does not have a recoating blade. Since this project was initially started, the university has acquired a Z-Corp machine whose accuracy is perfectly acceptable for the types of pre-surgical planning models used in this project. The making of the silicon rubber molds was a very time consuming and expensive process. In some cases, it was quite difficult to properly vent the molds, and the resulting polyurethane castings had trapped air pockets. While the flexibility of the rubber molds made it possible to demold complex shapes with small to moderate undercuts, demolding was still a very difficult proposition. Due to the need to produce the molds very quickly, one-piece cut molds were produced for all parts instead of two-piece molds that had built-up parting lines.

Due to the need to produce the molds very quickly, one-piece cut molds were produced for all parts instead of two-piece molds that had built-up parting lines. The process of cutting the molds with highly complex parting lines was a significant challenge. Final assembly of the separately cast bones was more difficult than it needed to be due to the fact that interlocking features were not added to the parts in the CAD system. The different parts were glued together using fast setting two- component epoxy glue with a cure time of 2 minutes. Keeping the bones aligned, and holding them in the right location for over two minutes while the glue cured was a challenge. In hind sight, it was decided that building parts in the QuickCast build style and then filling them with polyurethane would be much faster and less expensive than building RTV molds.

In hind sight, it was decided that building parts in the QuickCast build style and then filling them with polyurethane would be much faster and less expensive than building RTV molds. Furthermore, a positively locating joint between each bone will be designed to facilitate the assembly of future models.The results of the pre-surgical rehearsal and the surgical procedures were a success as well. Post-operative radiographs from both the first and the second surgery showed excellent limb alignment. The final surgical result will not be available until the end of August 2002 when the second external fixator will be removed. However the surgeon is confident that the result will be better than first anticipated and the dog will most likely be able to ambulate normally without any noticeable deformities. The dog ambulated with difficulty before surgery and was improved after the first limb was operated.

The dog ambulated with difficulty before surgery and was improved after the first limb was operated. It is anticipated that the second surgery will fully restore mobility of the pelvic limb.DiscussionUsing biomodels to plan and rehearse complex osteotomies of severely deformed bones has obvious benefits. The models give the surgeon a chance to determine if the planned procedure will succeed or if it is better left alone. Further, the ability to fully plan and rehearse a surgical22procedure in advance saves valuable time in the operating room, and substantial cost savings can be experienced. The time a patient is under anesthesia is critical and should be minimized. Routine cases would probably not benefit as much from the use of biomodels, but the more complex and unusual cases definitely will. Even cases earlier considered impossible can now be performed with help of new tools made available to the surgeons.

Even cases earlier considered impossible can now be performed with help of new tools made available to the surgeons. It could be argued whether or not it is economically and politically correct to spend money on biomodels to treat animals with severe deformities. It is noted, however, that many pet owners are very committed to their animals and would do almost anything to help them. The money spent on the biomodels will, in most cases, save enough time in the surgery room to cover the cost of the models. Further complications can also be avoided since the surgeon has more information to base his or her decision on, and the risk for making a mistake is therefore reduced. According to the surgeons involved with this project, they were much more confident when they started the surgery, since they had already done the entire procedure and had a very good idea of what to expect.

Models and experiments determine process parameters, including molding latitude, and ejection force requirements. Ejection force predictions are based on work by Menges, using values for elastic modulus determined from tensile tests at ejection temperatures. Similar studies are planned for stereolithography inserts (SL 5170).IntroductionManufacturers who currently build products in low volume, such as aerospace systems,can benefit from tools that will cost effectively produce low quantities of production parts. Injection molding, which is typically a very high volume process, requires significant decreases in tooling costs in order to make low quantity production feasible. The application of freeform fabrication techniques, such as laser sintering, to build injection mold inserts is one approach to reducing these tooling costs.Injection molds for high production volumes are traditionally machined of steel, are verystrong, and have good thermal properties.

The material properties of tools built using solid freeform fabrication (SFF) vary from conventional molds (see Table 1) but still may be suitable for injection molding lower quantities of parts. SFF processes are also attractive because they can generate complex geometries as easily as simple ones, e.g., they can build mold shapes and cooling lines that are impossible to machine. Selective Laser Sintering (SLS®) is a good example of a SFF process that can be used for injection mold inserts for low production volumes.Injection molding simulations and experiments will be run with thermoplastic materials,first using a machined steel mold insert and then using a SLS® insert (see Figure 1). The objective is to determine process parameters and ejection force requirements for the sintered insert and compare them to those for the steel insert.

The objective is to determine process parameters and ejection force requirements for the sintered insert and compare them to those for the steel insert. A modular injection mold having a steel Master Unit Die mold base will be used, the core and cavity of which can be removed and replaced with those of other materials. Machining allowances were included in the design of the inserts so that they can be machined to fit properly into the mold base. The chosen part for this92research is a vented closed-end cylinder, similar to the plastic canisters used to store 35 mm photographic film (see Figure 2).Table 1: Properties of steel and SFF mold materials.ProcessMold MaterialDensitykg/m3Tensile Strength MPaHardnessBaseline [1]MachiningP-20 Mold Steel7870108030-35 RCH-13 Tool Steel7800155052 RCRapid Tooling Materials *3D Printing – PrometalBronze/infiltrant810040660 RBLaser Sintering – 3D SystemsSteel, w/copper345033.675 ShoreDS.

Steel, w/bronze770051079 RB as machinedLaser Generating – LENSS. Steel 3168000 [2]80080 RB [2]Plastic Casting – CIBACeramic-filled Epoxy64 (UFS)91 ShoreDStereolithography – 3D SysSL 5170 cured resin12205985 ShoreDFrom company literatureFigure 1: SLS® core and cavity insert made with LaserForm™ ST-100.93ConductivityW/mC47.6 @ 204C25.1 @ 199C7.351.28 at 40C 0.92 at 150C49 at 100C 56 at 200C15 [2]0.200Figure 2: Canister part.TheoryProcess ParametersA typical manufacturing environment in which very large quantities of injection molded thermoplastic parts are produced requires high quality products and minimal cycle times. Since viscosity of the thermoplastic melt decreases with increasing shear rate (see Figure 3), injection velocities are kept as high as possible to allow the mold to fill quickly and completely.

Since viscosity of the thermoplastic melt decreases with increasing shear rate (see Figure 3), injection velocities are kept as high as possible to allow the mold to fill quickly and completely. Shorter cooling times are also favorable.Polystyrene1400) s - a P12001000(y t i800s o c s Vi60040020000100020003000400050006000Shear Rate (1/s)T=190CT=210CT=225CFigure 3: Viscosity vs. shear rate following the power-law model [3].For the production of small quantities, part quality is still very important, but there is less emphasis on minimizing cycle times. If cycle times can be relaxed, then injection velocity may be decreased and cooling times may be increased. This allows for core and cavity inserts of different materials, such as SFF materials. Cores made with some SFF materials may not be able to withstand the pressures and temperatures used with steel molds.

Cores made with some SFF materials may not be able to withstand the pressures and temperatures used with steel molds. A balance must be found between velocity and viscosity, so that the polymer fills the mold completely and produces a quality part.94SLS® material LaserForm™ ST-100 has half the strength of mold steel, but comparablethermal conductivity. It is expected that injection velocity and temperature will have to be changed to some extent compared to those of the baseline steel insert. A simulation of injection force using a sintered ST-100 insert (Figure 4) shows that the core can withstand a reasonable injection velocity. At an injection pressure of 109 MPa, maximum deflection of the core at the point of injection is minimal (0.003 mm). For comparison purposes, a similar model is shown for a stereolithography SL5170 core.

For comparison purposes, a similar model is shown for a stereolithography SL5170 core. For the same injection pressure, deflection of the core is significant (0.3 mm).Figure 4: A simulation of stress due to injection pressure on an ST-100 core (left) and an SL5170 core (right).Ejection ForceEjection forces have two primary components: opening forces and, more importantly in this case, release forces. The mold material, the part material, and the processing conditions are RF can be computed all factors affecting release forces. For sleeve-type parts, the release force from the coefficient of friction f , the contact pressure between the part and the core surface area of the coreAp , and theCA , as follows [4].F R×= fpA×A CFor cylindrical sleeves, the part shrinks onto the core, and stresses are subsequently builtup. Immediately upon ejection, the part recovers.

Immediately upon ejection, the part recovers. According to Menges, the contact pressure can be estimated from the contraction of the part diameter or circumference.

The relative change in diameter is given by ( td i d)−d∆ d∆==ceCffcwherefC∆ cd = core diameter ( i td= relative change in circumference)= inside diameter of sleeve immediately after ejection.e95Applying Hooke’s Law, σ × = E ε where σ = stressE = elastic modulusε = strain.Since, in this case, ∆=ε∆=dCffthen =σ( ) TE e ( eTE∆× )dfwhere= elastic modulus at ejection temperature.Contact pressure for the sleeve is given by )( TE e×∆×σ×dssfm==mpAr cr cwhere ms = wall thicknesscr = core radius.The surface area of the core is π L A C where L = core length.=××dcSo, with the coefficient of friction, release force is s)( TE e×∆×dπ L××=fm××fF Rdcr cNote that the coefficient of friction must be determined at process conditions duringejection, and modulus must be determined at ejection temperature.

In this work modulus was measured by tensile testing polymer specimens at temperature, and ejection force will be measured experimentally using load cells behind the ejector pins. Coefficient of friction will then be determined using Menges’ equation. Future release forces can therefore be calculated for these materials at these process conditions.MethodologySLS® ProcessThe laser sintering process used in this research involves a polymer-coated 420 stainless steel-based powder, known as LaserForm™ ST-100, and a 3D Systems Vanguard™ machine. Specifications of the Vanguard™ and material properties of ST-100 are shown in Tables 2 and 3.When the 3-dimensional part is initially built on the Vanguard™ System, the laser heatsthe metallic particles above the glass transition temperature of the polymer coating. The polymer softens and deforms, then fuses with other particles at each contact surface.

The polymer softens and deforms, then fuses with other particles at each contact surface. The temperature is96such that melting of the metal does not occur, only viscous flow of the polymer coating. The metal powder is then bound together by the polymer to form the “green” part. After the build is complete, the green part is removed from the Sinterstation and excess powder is brushed away. A furnace cycle follows in a reducing atmosphere to burn off the polymer, sinter the steel powder, and infiltrate the part with bronze. Infiltration eliminates any voids within the steel, resulting in a fully dense part. [5][6]Table 2: Vanguard™ System specifications (3D Systems).Model Number Laser Wavelength Power Beam Diameter Max. Scan Speed Min.

[5][6]Table 2: Vanguard™ System specifications (3D Systems).Model Number Laser Wavelength Power Beam Diameter Max. Scan Speed Min. Layer Thickness Build ChamberLC-100 DEOS CO2 Laser 10.6 microns 100W max at part bed 450 microns 10,000 mm/sec (394 in/sec) (0.10 mm) 0.004 in 381w x 330d x 457h mm (15w x 13d x 18h in)Table 3: LaserForm™ ST-100 material properties (3D Systems).Density Thermal ConductivityCTE Tensile Yield Str. (0.2%) Tensile Strength Young’s Modulus Elongation Compression Yld Str (0.2%) Hardness, Rockwell B7.7 g/cm3 49 W/moK @100oC 56 W/moK @200oC 12.4 ppm/oC 305 MPa 510 MPa 137 GPa 10% 317 MPa 87 As infiltrated 79 As machinedASTM D792 ASTM E457 ASTM E457 ASTM E831 ASTM E8 ASTM E8 ASTM E8 ASTM E8 ASTM E9 ASTM E18 ASTM E18Modeling and SimulationMold fill simulations using MoldFlow® provide some validation of the mold design and predict what some of the processing parameters might be for the steel insert.

The values for maximum contact pressure will be used with friction coefficient and surface area in the Menges equation to calculate required ejection force. Results of such simulations will be compared to experimental data, and a model of ejection force will subsequently be created.97Figure 5: MoldFlow® image showing canister mold fill time for a steel insert.Figure 6: Contact pressure of the canister on a SL 5170 core prior to (left) and during(right) ejection.Tensile TestsElastic moduli for high density polyethylene (HDPE) and high impact polystyrene(HIPS) used in this research were measured at various temperatures using ASTM D 638 “Standard Test Method for Tensile Properties of Plastics” as a guide. The testing apparatus was an Instron model 1322 tensile tester with a tube furnace.

The testing apparatus was an Instron model 1322 tensile tester with a tube furnace. An extensometer with a 2-inch gauge and 50 percent strain was used to measure elongation.ASTM Type I (dogbone) specimens of each thermoplastic material were pulled at room temperature and at ten degree increments, starting at 30oC, until no elastic region was detected. HDPE was tested up through 70oC, and HIPS was tested up through 60oC. Results are shown in Figure 7. The values for elastic modulus will be used in ejection force calculations.ExperimentsInjection molding experiments will be run on a Sumitomo General Injection Molding Machine, model SH50M, a horizontal press with a fully hydraulic, 50-ton clamping system.98A series of injections with varying velocity and temperature will be run for each insert material to determine suitable processing windows. Values for the ejection force as measured by the load cells and canister diameter immediately after ejection will be recorded.

Values for the ejection force as measured by the load cells and canister diameter immediately after ejection will be recorded. Using Menges’ ejection force equation and the elastic moduli described above, values for coefficient of friction will be determined for each core material, and the validity of the equation will be checked.Modulus vs. Temperature350030002500HDPE HIPSa P M s u u d o Ml200015001000500020304050607080Tempe rature deg CFigure 7: Modulus vs. Temperature for HDPE and HIPS.SummaryIf tooling costs can be greatly decreased, injection molding becomes a viable process for production of small quantities of parts. One approach to this is the use of SFF processes for making injection mold inserts. The research described in this paper will provide useful data on the feasibility of this rapid tooling approach.

plication of Stereolithography in the Fabrication of Rehabilitation AidsSuresh J ayanthi, Dr. William Harwin 1, Dr. Michael Keefe 2 University of Delaware / A.L duPont Institute P.O.Box 269, Wilmington, DE 19899 (302) 651-6830, email: jayanthi@me.udel.eduDr. Vijay Kumar General Robotics and Active Sensory Perception Laboratory School of Engineering and Applied Science University of Pennsylvania Philadelphia, PA 19104.AbstractFree form fabrication methods have a great potential to signifi cantly improve the design and manufacture of equipment for people with physical disabilities, such as quadriplegia through spinal cord in jury, arthrogryposis, or cerebral palsy. Depending on the nature of the disability a device may need to be designed or adapted. A person with quadriplegia, for example, may benefit from an assistive device that maps existing head movements, into the movements of a spoon between a plate and his/her mouth.

A person with quadriplegia, for example, may benefit from an assistive device that maps existing head movements, into the movements of a spoon between a plate and his/her mouth. To be comfortable and effective a person Iuay need a headband that can connect to a suitable linkage. Stereolithography lends itself well to the fabrication of such one-of-a kind devices. Since the fabrication process time is considerably less than conventional approaches a greater number of iterations can be performed during the design to arrive at the most compatible device for a particular disability. Some rehabilitation devices are fabricated to establish the viability and linlitations of this approach.

Some rehabilitation devices are fabricated to establish the viability and linlitations of this approach. This paper details the efforts underway to integrate stereolithography with the needs of rehabilitation engineering.1Applied Sciences and Engineering Laboratory 2Mechanical Engineering Department2701IntroductionRehabilitation (or assistive) devices require diverse approaches to design and manufacture owing to the fact that their requirements are often highly spe cialized. The range of disabilities and the effects of those disabilities is ex tremely large. Their design is often complicated because of the intilnate interface between the device and the human body [1]. A number of complex biological variables need to be addressed and most often it is not possible to cOlnpile a thorough specification for a particular aid. Unfortunately, the number of people for whom the same device is suitable is small.

Unfortunately, the number of people for whom the same device is suitable is small. Though there are lllany fabrication techniques that are suitable for large volume produc tion, Inany of these are based on the fabrication of a mold from which parts are formed [2]. Depending on the complexity of the part the cost of making a Inold varies significantly. Since these molds are usually very costly it is not possible to include any Inodifications to design until the original cost of the first mold is recovered. However, with the advent of a whole new class of Inanufacturing approaches called free from fabrication (FFF) it has now becolne possible to custolllize the design of a rehabilitation aid to suit the specific requirements of a particular disability.

Since these FFF methods produce parts considerably faster than conventional approaches, the cost and cycle tilne are significantly lower.An alternative methodology for design and fabrication of rehabilitation devices has been proposed which attempts to separate the design of the user interface from the design of supporting features like wheelchairs, large but ton telephones and easy grip cutlery [1]. This Inethodology is based on the underlying assulllption that Inost of these supporting features can be read ily obtained from traditional techniques while the user interface needs to be specifically designed for a particular cirCUlnstance. Some of the exalllpies where this methodology is quite useful include wheelchair seating and con It is in the fabrication of trols, prosthetic anns and legs, and Magpie [4]. these user interfaces that the FFF techniques become highly desirable.

these user interfaces that the FFF techniques become highly desirable. This study is focused particularly on stereolithography, which is one of the more highly developed FFF techniques.271Data AcquisitionDiagnosisConsumer FeedbackPhysical PrototypingFigure 1: Design and rapid prototyping of one-of-a-kind assistive devices in rehabilitation2 Design MethodologyA rehabilitation device that can assist in tasks of daily living need not be cOlnplex or have a universal function to be successful. Simple designs have the attraction of increased reliability, lower costs and a more ready accep two of the most widely used rehabilitation devices namely, tance. Handy II [3] and the Magpie [4] are quite simple in design, perform many tasks and are suitable for a wide range of disabilities. However, the main dis advantages of these devices are lack of provision for significant adaptation, long manufacturing time and high cost of the product.

However, the main dis advantages of these devices are lack of provision for significant adaptation, long manufacturing time and high cost of the product. The design of reha bilitation equipment requires the quantitative assessment and measurement of the disabled person's perfonnance and a methodology to quickly design and prototype assistive devices based on performance goals. The approach adopted for the fabrication of aids is illustrated in Figure 1. The approachIn fact272involves,.. a quantitative estimation of the form and performance requirements of the assistive device based on the specific requirements of a disabled person,..the design of the assitive device,.. evaluation of the device using simulation and virtual prototyping,.. feedback from the consumer and the therapist or physician,.. actual prototyping of the assistive device using free form fabricationmethods,.. evaluation of the function and performance of the device and,..

actual prototyping of the assistive device using free form fabricationmethods,.. evaluation of the function and performance of the device and,.. redesign based on the performance.The basic objective here is to quantitatively assess the capabilities of a dis abled individual which will in turn lead to the design problem specifications. For example, a person with a high level spinal cord injury may have limited shoulder and elbow movement with no ability to grasp objects in the environ ment and limited capabilities to supply adequate force to a telemanipulator. We can characterize his or her range of movements and the torques/forces that are applied during these movements.

We can characterize his or her range of movements and the torques/forces that are applied during these movements. From this information a telethesis linkage can be designed to fit to the person - one that will harness his or her movements and force capabilities in an optimal fashion.3 DiscussionThe eventual goal of this study is to fully integrate the fields of rehabilitation engineering and free form fabrication through the design methodology de scribed above. One of the primary advantages of this approach is the ability to involve the end-user in the design process. Since the end-user of the device is the person with disability, the design optimization is essentially based on273Data Acquisition ~ Data ProcessingData Conversion .....,rPhysical Prototyping....Stereolithography .- .....Solid ModelingFigure 2: Path from Data Acquisition to Physical Prototypethe feedback from that person.

This paper establishes the feasibility of the approach by actually fabricating specific assistive devices. Stereolithography is the FFF technique adopted in this study. Two devices namely,the head band and the mouth-stick are presented here. The head-band is designed to be part of a head-controlled telethesis for feeding a person with disabil ity. Such a head-controlled feeding device is especially useful for people with quadriplegia. 3 These serve as two generic examples to validate the proposed design methodology.Figure 2 shows the various steps involved in the process, starting from data acquisition to generation of the physical prototype. To begin with, the image of the head was obtained by scanning the head froIn different angles and registering the different images. Subsequently, the image is subjected to a series of data filtering algorithms in order to extract a rectangular strip of the image along the contour of scan, passing through the forehead.

Subsequently, the image is subjected to a series of data filtering algorithms in order to extract a rectangular strip of the image along the contour of scan, passing through the forehead. The image thus obtained is coverted to a neutral format such as IGES using cus tomized software and later filtering algorithllls are used to refine image. The extracted surface feature is later imported into Pro-engineer and ex-3When quadriplegia is a result of a spinal cord injury, depending on where the lesion occurs the upper extremities and even the torso may be impaired. But the quadriplegic generally has good control over his/her head and neck movements.274Figure 3: Profile of the Head-Bandtruded to created a solid strip of uniform. cross-section. However, in order to overCOlne the inherent unevenness built into the surface, an elliptical cur vature was fitted to closely follow the contour of the scanned image.

cross-section. However, in order to overCOlne the inherent unevenness built into the surface, an elliptical cur vature was fitted to closely follow the contour of the scanned image. This elliptical feature was later extruded to create a solid model of the head- band. In order to fabricate this solid model a tesselated representation (.STL for formatted representation of the solid model is mat) is generated. This then used to prepare the part for fabrication in an SLA machine. Figure 3 illustrates the contour of the head-band obtained from interpolating a ellip tical curvature to a range scanner iInage of a subject's head obtained using a Cyberware 3030PS scanner. Figure 4 shows the geollletry of the mouth stick which was obtained by remodeling and modifying a customized design. The cad model thus developed will be used prototype the mouth-stick in a SLA machine.

The cad model thus developed will be used prototype the mouth-stick in a SLA machine. The main thrust here is to develop a simple device which ad dresses the routine tasks like lifting sheets of paper on a table. The slit seen in Inouth-stick is intended to meet this purpose. A series of iterations will be performed to establish a design loop that produces the mouth-stick best suited for the person's needs. The next step in the development of this device is the interface that rests against the upper jaw of the person and allows easy275Figure 4: Profile of the Mouth-Stickmanipulation. Here, to ensure that the interface is non intrusive the contour of the upper jaw needs to be scanned and the interface developed to rest comfortably against the jaw bone.

Here, to ensure that the interface is non intrusive the contour of the upper jaw needs to be scanned and the interface developed to rest comfortably against the jaw bone. The eventual goal of this whole approach is to arrive at the most optilllal design solutions and to develop algorithllls which will streamline the whole design phase into a series of routine steps.4 Future DirectionCurrently the work is focused on specific assistive devices with a view to establishing the feasibility of this approach in the design and fabrication of such devices using the free form fabrication approach.

Eventually, however, the goals of this research are the following.(I The developlllent of new cOlllputer-integrated design tools for the quan titative asseSSlllent of function and performance requirements of assis tive devices.(I The integration of all relevant tools and technologies into a practicalrapid prototyping environlllent.276The development of mechanisms for the systematic evaluation of the final product.5 ConclusionsThough the stereolithography approach to fabricate rehabilitation devices is quite efficient it is yet to beconle cost effective. The high cost involved in employing this technology is a major bottleneck in the widespread use of this approach. Another major difficulty that needs to be overcome is the toxicity and other material related issues involved in the use of photopolymers in contact with the human body for prolonged periods of tinle.

A lot of rehabilitation devices involve fabrication of small or even tiny parts and hence there is need for accurate desktop machines which could quickly be used to verify new design formulations. With continuous improvements in the Inachine accuracy as well as ever increasing class of photopolymer materials it is hoped that the approach outlined in this study will eventually become widely used in the rehabilitation industry.References[1] Roger D. Orpwood, Design Methodology for Aids for the Disabled, Journal of Medical Engineering and Technology, vol. 14, p2-10, 1990.[2] Albert M. Cook, Assistive Technologies: Principles and Practice, Mosby1995.[3] John R. Hegarty and Mike J.

14, p2-10, 1990.[2] Albert M. Cook, Assistive Technologies: Principles and Practice, Mosby1995.[3] John R. Hegarty and Mike J. Topping, HANDY1 - a low cost robotic aid to eating, 2nd International Conference on Rehabilitation Robotics, p17-25, Georgia Institute of Technology, Bioengineering Research Center, Atlanta, Georgia, June 1991.[4] Mervin Evans, Magpie - It's Development and Evaluation: Internal Re port, Oxford Orthopeadic Engineering Center, Nuffield Orthopeadic Cen ter, Headington, Oxford, England OX3 7LD, 1991.277.

viewed PaperApplication of Systems Engineering to Lattice Structure Design: A Framework for Requirements Flow DownJoseph W. Fisher1,3, Timothy W. Simpson1,2, Simon W. Miller31Department of Mechanical Engineering, The Pennsylvania State University, University Park, PA 16802 2Department of Industrial and Manufacturing Engineering, The Pennsylvania State University, University Park, PA 16082 3Applied Research Lab, The Pennsylvania State University, State College, PA 16803AbstractIn the field of additive manufacturing, highly ordered cellular structures with repeating patterns in space, known as lattice structures or simply “lattices”, have been shown to sig- nificantly improve the performance per unit mass of engineered components compared to traditionally machined designs. The characteristics of these lattices are dependent on many design decisions, leading to challenges in determining how to best utilize them.

The characteristics of these lattices are dependent on many design decisions, leading to challenges in determining how to best utilize them. This issue is further complicated when the density, size, and topology of the lattice can spatially vary in the design domain. Many methods have been proposed to optimally distribute material within a lattice structure; however, these methods exist in a bubble, independent of broader engineering design criteria and processes. In this paper, we introduce a framework based on systems engineering to guide the design of lattice structures using the component requirements as inputs and outputting a verified and validated design based on those requirements. The framework translates component requirements down to the lattice structure and provides a means to verify and validate that the final design meets those requirements.

The framework translates component requirements down to the lattice structure and provides a means to verify and validate that the final design meets those requirements. We provide a walk-through of the framework by designing a lattice structure for an oil and gas component, and discuss future research to extend, verify, and validate the framework.1 INTRODUCTIONAdditive Manufacturing (AM) allows for the realization of greater freedom and complexity in de- sign by offering access to previously untapped design space during the manufacturing process [1]. One type of complexity that is enabled by AM is hierarchical complexity where the features of the geometry are composed of one or more levels of smaller features [2]. This hierarchy can be achieved by incorporating cellular solids into the component [3].

This hierarchy can be achieved by incorporating cellular solids into the component [3]. Cellular solids are hierarchical materials that contain two or more phases; they are found both in natural materials such as wood and bone, as well as engineered materials including foams [4]. AM has enabled the creation of cellular solids with much greater control over their geometry compared to traditional engineered foams [5]. Highly-ordered cellular solids with repeating geometric patterns are known as lattice structures or simply lattices [6]. Integrating lattices into engineering designs offers the potential for significant reduction in weight compared to those designed for traditional manufacturing [7].

Integrating lattices into engineering designs offers the potential for significant reduction in weight compared to those designed for traditional manufacturing [7]. In2174addition to reducing the amount of material consumed by the AM process, lattices can also reduce manufacturing time as the volume of deposited material directly correlates to manufacturing time for most AM processes, both of which directly relate to the total cost of production [8].Lattice structures contain repeating patterns and can be generated by patterning a small portion of the structure in space. The smallest portion of the structure that can be patterned to produce a given lattice is known as its unit cell. The unit cell is often used in discussion and comparison of different lattice typologies.

The unit cell is often used in discussion and comparison of different lattice typologies. Unit cells can also be used to compute the homogenized properties of the lattice because of the repeating nature of the structure [9].The hierarchical complexity of AM lattices offers unique challenges related to modeling, analysis, and inspection of the lattices due to the significant number of small features relative to the size of the whole component – this leads to a large number of elements needed to accurately represent the features of the lattice for manufacturing, or even more so, during analysis using finite element methods [10]. These three tasks become even more complex when the lattice is non-uniform (spatially varying) or optimized so that each cell of the lattice is tailored to meet the lattice level requirements of the application [11, 12, 13].

The number variables that must be controlled for effective lattice design is a potential reason why lattices have not seen wider adoption in application. Advances in software tools makes the modeling and analysis of lattice structures more accessible [14, 15, 16]. However, this still leaves engineers with a larger number of design decisions that must be made as well as the need to analyze and inspect the complex internal features [17] to verify its required performance.The unique opportunities and limitations related to AM has given rise to the concept of Design for Additive Manufacturing (DfAM) [18]. Several frameworks or approaches based on DfAM principles and methods have been proposed, both broadly and tailored to specific DfAM opportunities including lattice structure design. For instance, Bikas et al.

For instance, Bikas et al. [19] noted the importance of consideration of DfAM principles early in the design process and proposes a framework that propagates AM process information back through the process to eliminate the need for iteration. Chennamkulam [20] developed a framework using inverse problem solving and a TRIZ-based method methodology to redesign components with DfAM considerations. Li et al. [21] proposed lattice optimization using homogenization to map between the properties and the volume fraction (VF) of a lattice. Y. Wang et al. [22] used topology optimization technique to optimize lattice structures, and C. Wang et al. [23] developed a two-step method of lattice optimization that first uses topology optimization, followed by sizing optimization of the elements of the lattice. Gorgu- luarslan et al.

Wang et al. [23] developed a two-step method of lattice optimization that first uses topology optimization, followed by sizing optimization of the elements of the lattice. Gorgu- luarslan et al. [24] proposed a two-step optimization first using ground structure optimization, with a secondary step that accounts for limitations of the additive process to produce manufacturable designs. Chen et al. [25] proposed a lattice generation and optimization method using finite element meshing to define the cells of the lattice and the MIST topology optimization scheme to optimize the cells [26]. Vaissier et al. [27] investigated lattice unit cell and parameter selection for static heat dissipation.These frameworks provide the basis for designing with lattices but they lack direction and rigor when integrating components with lattices into a system – a problem that we address in our work herein.

Moreover, existing lattice design frameworks do not incorporate considerations2175for the selection of unit cell(s) based on the local loading; these decisions are therefore often made without explicit consideration of the performance required of the lattice. It is challenging to draw direct comparisons between different lattice structures because of the variance between studies that publish data (e.g., material properties, lattice parameters, process parameters, sample replicates, testing standards). However, by delaying the unit cell selection until performance requirements have been established, informed decisions can be made, allowing for more optimal material utilization [11]. Recent work to bring together and normalize data from the literature has offered a better understanding of the relative performance of different lattice structures [28].

Recent work to bring together and normalize data from the literature has offered a better understanding of the relative performance of different lattice structures [28]. A database of strut-based unit cells was developed and used in an optimization scheme that selects the topology of each cell in the lattice from the database [29]; however, determining the independent effects of most lattice parameters on the material properties is still difficult based on current data.When it comes to translating requirements from a system to a subsystem or component, several models exist in the literature including the Waterfall [30], Spiral [31, 32], and System Vee [33] models. The System Vee was developed to resolve perceived short comings in prior models, including defining the role of systems engineering in the process [33].

The System Vee was developed to resolve perceived short comings in prior models, including defining the role of systems engineering in the process [33]. The System Vee has been widely adopted into the System Engineering Process [34, 35]; however, multiple definitions of what the model encompasses and the steps involved exist [33, 35, 36, 37], as is the case with most of these methods.The System Vee does highlight and accommodate the hierarchical structure of a system that is composed of subsystems and components. Most extensions and adaptations of the System Vee share the same starting point at the top left of the Vee which is the definition of user requirements for the system. These requirements flow down the left side in a process that mirrors the Waterfall model, and on the right side, incremental steps build up to the full system with verification and validation (V&V) [33].

Models at the system level, including the System Vee, do not specify the method for detail design of the subsystems and components of the system, leaving this to the design engineers to select appropriate processes.Currently, there is no framework for the design of lattice structures within a requirement- driven engineering process. The development of such a framework would enable lattice structures within the design of complex engineering systems. The major contribution of this work to the literature is the development of an accessible, systematic, and extensible framework for using component requirements to inform the design, integration, and V&V of lattice structures. This framework serves as a single engineering process that enables engineers to more readily incor- porate lattice design into system design. In Section 2 we introduce the Lattice Vee Framework and its steps along with the assumptions necessary to use the framework.

In Section 2 we introduce the Lattice Vee Framework and its steps along with the assumptions necessary to use the framework. To demonstrate the framework, we walk-through with an example component in Section 3.

To demonstrate the framework, we walk-through with an example component in Section 3. Based on this presentation and demonstration of the Lattice Vee, we discuss the implications and future research directions in Section 4, and end with closing remarks in Section 5.2 FRAMEWORK BASIS AND DEVELOPMENTThis framework uses hierarchical decomposition drawing parallels between the hierarchical struc- ture of a system composed of subsystems and components, to a component with an integrated2176ComponentV&V PlanningLattice Level RequirementsComponent RequirementsComponent IntegrationComponent AnalysisLattice IntegrationLattice Cell VerificationUnit Cell ExplorationLatticeCell DefinitionLattice Domain DefinitionLattice Cell ManufacturabilityAM Process LimitationsLattice Verification PlanningFigure 1: Lattice Vee Frameworklattice structure composed of solid regions and lattice domains that can be broken down further to the individual cells that make up the lattice.

In both cases, the performance at the top-level is dependent on the performance of the elements of which it is built. Following systems thinking, it is then possible to follow this process in reverse in our framework and flow down requirements from the component to the cells of the lattice, allowing the requirements to drive design at the next level down in the hierarchy.While the System Vee model is primarily a descriptive model without explicit actions, the Lattice Vee Framework is more prescriptive, dictating specific actions at some steps while intentionally omitting others to allow for different methods of defining lattice level requirements, optimization techniques, and levels of complexity in the lattice design. This allows the Lattice Vee to be extensible while providing guidance on specific steps and actions to take. The framework is presented graphically in Fig.

This allows the Lattice Vee to be extensible while providing guidance on specific steps and actions to take. The framework is presented graphically in Fig. 1.Like the System Vee, moving down the left side of the Lattice Vee moves down a hierarchy, decomposing the component requirements into lattice level requirements that spatially vary across the component domain. At the bottom of the Vee, the detailed design decisions relating to the lattice geometry are made accommodating process limitations and the lattice level requirements. Moving up the right side, the lattice is integrated, verified, and validated against its requirements. Through this process, the Lattice Vee offers a requirement-driven, quantitative means of selecting lattice unit cells and optimizing lattice parameters.As shown in Fig. 1, the Lattice Vee starts with the component requirements, which should be derived from the requirements one level up in the system hierarchy.

In addition to component requirements, the Lattice Vee relies on three pragmatic assumptions for its use.2177Material and AM process are given, or have been selected.Component domain is of an appropriate scale for printing with the chosen process/machine.Database of sufficient lattice structure properties exists (current limitations discussed in Sections 3.8 and 4).Note that these assumptions are not intended to and should not limit design freedom; rather, they are necessary to provide information to help traverse and implement the framework.After establishing the (1) Component Requirements and (2) AM Process Limitations, and after satisfying the assumptions of the Lattice Vee, the following steps of the framework are performed:(3) Component Verification & Validation Planning, where V&V tasks are defined relating to thecomponent and manufacturing process;(4) Component Analysis, where the component requirements are used to inform analysis of thecomponent;(5) Lattice Level Requirements, where the component analysis is used to define spatially varyingmaterial property requirements;(6) Lattice Verification Planning, where verification tasks are defined for the performance of thelattice;(7) Lattice Domain Definition where the lattice sub domain of the component is defined;(8) Unit Cell Exploration, where the lattice unit cell(s) are selected using data driven decisions;(9) Lattice Cell Definition, where the selected unit cells are implemented within the lattice do-main with properties based on the lattice level requirements;(10) Lattice Cell Manufacturability, where the cells of the lattice are tested to identify potentialissues in manufacturing;(11) Lattice Cell Verification, where the verification of the performance of different cells withinthe lattice takes place;(12) Lattice Integration, where the cells of the lattice are joined and geometric discontinuitieswithin the lattice are resolved; and(13) Component integration, where the lattice joins with the rest of the component and V&V canbe carried out.2178The numbering of this list corresponds with the numbering of the following subsections, where each of these steps of the Lattice Vee Framework are discussed in more detail, as well as the numbering in Section 3 which contains a walk-thought of the framework.2.1 Component RequirementsThe requirements for the component flow down from the system level requirements.

They serve as the starting point for this framework. Component requirements must include the design domain of the component and performance specifications for the component. This provides the framework the spatial volume in which it can operate, and defines what is trying to be achieved by incorporating lattice structures.2.2 AM Process LimitationsAll AM processes have a unique set of process limitations (e.g. allowable overhang angles, minimum feature sizes, and support removal requirements) [19, 38, 39] that are also effected by the machine, material, and the processing parameters used [39]. In this framework, these limitations are directly applied during the four steps located at the bottom of the Vee that are highlighted in Fig. 1 and discussed in Sections 2.7 to 2.10.2.3 Component Verification & Validation PlanningBased on the component requirements, a plan for the V&V of the component prior to system integration must be defined.

For most AM components, this plan should include production, Non-Destructive Evaluation (NDE), post-processing steps (e.g., heat treatment and machining), material characterization, and physical testing. In addition to these required steps, the plan should include any other performance evaluation needed to ensure that the component requirements are met, allowing for integration into the system design.2.4 Component AnalysisRequirements of the component are used to inform boundary conditions to be used during sim- ulation and analysis. This simulation will produce a metric or set of metrics for performance of material within the component domain that varies spatially. We do not specify metrics or a method for defining these metrics directly to allow for adaptability to different situations and applications.

We do not specify metrics or a method for defining these metrics directly to allow for adaptability to different situations and applications. Metrics could be derived from a wide array of different means including: topology optimization, stress, strain or strain energy, buckling, fatigue, fluid interaction, heat transfer, thermal expansion, or any other relevant quantifiable measure [12, 22, 40, 41, 42, 43].2.5 Lattice Level RequirementsFrom the result of the component analysis and the bulk material properties, we can derive the requirements that apply to the material as a function of spatial location.

Since the cells of the lattice have not yet been defined, the lattice level requirements are not specific requirements for each cell of the lattice as the boundaries of each cell have not yet been defined, but rather, they are the basis on which the cells will be verified once defined.21792.6 Lattice Verification PlanningFor the verification of the lattice level performance of the lattice structure, we suggest homog- enization of the cells into constitutive relationships. Those homogenized properties can then be compared to the lattice level requirements within the volume of that cell.

Those homogenized properties can then be compared to the lattice level requirements within the volume of that cell. This verification can be applied to all cells or a representative set of cells (based on geometric similarity) to reduce the computational load and increase the speed of iteration and exploration.Note, we can only compare the performance of the lattice to metrics derived from the component requirements, it is not possible to validate the lattice structure prior to integration with the rest of the component, leaving validation to take place during the component integration (Section 3.13) as planned in Section 3.3.2.7 Lattice Domain DefinitionThe lattice domain is the space that remains when the solid regions of the component domain are removed.

These solid regions come from three factors: (1) functional surfaces (with thickness) or volumes, (2) lattice level required properties (relative to material properties), and (3) limitation based on what size unit cells, and lattice features, the process and material are capable of producing. The space that remains after these factors are taken into consideration is the lattice domain, which is the volume (or volumes) that the lattice will fill within the component.2.8 Unit Cell ExplorationThe definition of the lattice level requirements and the lattice domain constrains the range of properties required from the lattice structure. Based on this range, this step of the framework calls for a selection of a unit cell, or multiple unit cells that will fill the lattice domain.

Based on this range, this step of the framework calls for a selection of a unit cell, or multiple unit cells that will fill the lattice domain. By leaving the unit cell selection until this point in the process, intuition-based selection can be avoided and the decision can be data driven using a database of lattice properties [28].2.9 Lattice Cell DefinitionEach cell of the lattice must be defined to fill the lattice domain with structure. This includes definition of the size, shape, orientation, and position of each cell, as well as the amount of material in each cell, and how material is distributed within the cell — all of which can vary spatially throughout the lattice domain.The definition of these different parameters need not happen in this order or serially. As discussed in Section 1, there exist a number of different methods for defining and optimizing a lattice.

As discussed in Section 1, there exist a number of different methods for defining and optimizing a lattice. Therefore, the intent of this framework is not to supplant existing tools, but to offer a means to connect design and optimization of lattices to a requirement-driven design framework and analysis toolchain. To this end, we do not explicitly describe the pathway to defining the lattice parameters.21802.10 Lattice Cell ManufacturabilityThe unit cells typically used in lattices are generally considered manufacturable with common AM processes; however, the previous step leaves open the potential for significant variance within lattice structure that could lead to significant changes of the geometry. This creates the potential for cells that violate the specific manufacturability limits of the selected process and material combination, even when the original unit cell was manufacturable with the chosen AM process.

In this step, the manufacturability of each geometrically unique cell, or a set of cells that adequately represents the different cells of the lattice, should be tested for problems relating to: minimum feature resolution (on material as well as voids), bridging, self-supporting angles, trapped feed stock or material, islands, and pinched off features as well as any other manufacturing or process constraints [44, 45].2.11 Lattice Cell VerificationAs with the manufacturability test in Section 2.10, the homogenization routine used for verification needs to be carried out for each geometrically unique cell, or a representative set of cells. A comparison of the homogenized properties of each unit cell to the lattice level requirements that fall within that cell (based on the plan developed in Section 2.6).

A comparison of the homogenized properties of each unit cell to the lattice level requirements that fall within that cell (based on the plan developed in Section 2.6). This offers a verification of performance without the need to first integrate, mesh, and analyze the full-scale model of the complex lattice structure, which can be computational expensive [10].2.12 Lattice IntegrationBy treating cells of the lattice independently, there is increased potential for sharp corners and transitions that could lead to stress concentrations. Smoothing operations should be applied to mitigate these issues. Although this will change the properties of the cells from the homogenization performed in Section 2.11, this can be done conservatively to avoid weakening the structure and maintain manufacturability.

If the modeling technique produces separate geometric bodies within the lattice, then they would also be merged at this step to form a single, manifold geometry.2.13 Component IntegrationWith the lattice cells modeled, integrated, and smoothed, the overall lattice structure can be inte- grated with the rest of the component domain, and the component V&V plan (see Section 2.3) can be carried out. After this step, the Lattice Vee has been implemented, and a manufacturable design that meets the component requirements has been produced, verified, and validated.2.14 Iteration within the FrameworkAs with all design processes, iteration within the Lattice Vee may be necessary if the design fails to meet any requirements during the integration of the design. Such iteration could start with any step after “Lattice Cell Definition” leading back as far as “Lattice Domain Definition”. For visual clarity, the Lattice Vee shown in Fig. 1 does not show iteration loops within the framework.

For visual clarity, the Lattice Vee shown in Fig. 1 does not show iteration loops within the framework. This type of iteration within the framework marks an additional departure from the System Vee model.2181Table 1: Material properties for Inconel 718 taken from [46]Young’s Modulus Yield Stress Density28800 ksi 125 ksi 0.289 lbm/in3199 GPa 862 MPa 8000 kg/m3In the System Vee iteration is limited within the process by the control/phase gates used for project management [33].In addition, the system level design process is not linear and changes to the system level requirements result in changes to the component requirements. This often necessitates returning to the first step of the Lattice Vee.

This often necessitates returning to the first step of the Lattice Vee. Application of the Lattice Vee framework is discussed next for an oil and gas component with a uniform lattice.3 WALK-THROUGH OF FRAMEWORKTo demonstrate the application of the Lattice Vee, we present this walk-through of the framework applied to a component used in high pressure applications in the oil and gas industry. This component has been previously demonstrated to be a good candidate for lightweighting with lattice structures [46]. The component has large regions of material that do not significantly contribute to its functionality. Removing this material with a subtractive process would be both time-consuming and wasteful. Meanwhile, the internal passageways that give the part its functionality are difficult to machine, requiring multiple set-ups and operations.

Meanwhile, the internal passageways that give the part its functionality are difficult to machine, requiring multiple set-ups and operations. With AM, these internal channels are much easier to produce, and the weight of the part could be reduced simultaneously, which would reduce material and production cost.3.1 Component RequirementsThe domain of the component is a cylinder that is 127 mm long with a diameter of 76.2 mm and two non-intersecting internal fluid channels (see Fig. 2). The component must maintain a minimum Factor of Safety of 1.5 under the loads shown in Fig. 2. We take the minimum material thicknesses for the pressurized surfaces from [46] to allow for direct comparison of the lattices. For internal features, the minimum thickness is 5.08 mm, and the minimum thickness for the external cylindrical surface is 7.62 mm. The component is to be manufactured out of Inconel 718 using an EOS M280 Laser Powder Bed Fusion (LPBF) system.

The component is to be manufactured out of Inconel 718 using an EOS M280 Laser Powder Bed Fusion (LPBF) system. We present the material properties used in Table 1. We constrain the cells of the implemented lattice to be uniform for ease of inspection and to mirror the original design.3.2 AM Process LimitationsBased on the material and process that were specified in the component requirements (see Sec- tion 3.1), there are limitations on feature size, unsupported overhanging angle, and bridging dis- tance that impact the design of the lattice structure.

For the purposes of this application, we define the limiting values as: 0.5 mm, 40◦, and 5 mm, respectively.2182Figure 2: Primary dimensions and loads applied to internal channels and outer cylindrical surface21833.3 Component Verification & Validation PlanningAt this point, the details of the design have not yet been defined, however, we are able to plan for the verification and validation of the design based on the component requirements. For the verification of this design, we plan to first perform a finite element analysis of the component with the final selected lattice integrated. After the component is built using LPBF and stress relieved, micro-CT scanning will be used for NDE of the surfaces of internal fluid channels and the lattice, as well as inspecting for porosity throughout the component domain.

The component should then be heat treated and machined, after which the Geometric Dimensions and Tolerances (GD&T) would be evaluated, and the component would be pressure tested to ensure that the requirements are satisfied.3.4 Component AnalysisBased on the component requirements, we define a static structural finite element model using the given load conditions, the component domain (see Fig. 2), and material properties (see Table 1). From this we obtain the von Mises stress, shown in Fig. 3a, throughout the component domain to inform the lattice level requirements.3.5 Lattice Level RequirementsTo determine the lattice level requirements of the lattice, we divide the yield stress of the material (see Table 1) by the required factor of safety to determine the allowable stress of 574 MPa. We then normalize the von Mises stress to the allowable stress.

We then normalize the von Mises stress to the allowable stress. The result of this is an estimation of what fraction of the bulk material is needed spatially within the component to meet the stress requirement. The peak stress within the component domain corresponds to ∼6% of the allowable stress, indicating that significantly less material is needed to support the load. The peak stress falls within the solid sections defined in the component requirements. Once the lattice domain is established (in Section 3.7), we can find the peak stress within that sub-domain.3.6 Lattice Verification PlanningWhen analyzing the cells of the lattice structure, a single homogenization routine is needed since a uniform cell size and density are being used.

We can then compare the homogenized properties to the maximum of the lattice level requirements that falls within the lattice domain (see Section 3.7).3.7 Lattice Domain DefinitionAs discussed in Section 2.7, there are three sources that we check to define the regions of the com- ponent that must be solid. From the component requirements, we have solid regions at the interior faces, and external cylindrical face. From the component analysis, we find no additional solid material is needed beyond what is already given. This makes intuitive sense, because we based the surface thickness on the solid domain of a prior study that used the same requirements [46]. Lastly, there are no small features on the remaining volume that necessitate exclusion from the lattice domain. The result, shown in purple in Fig.

Lastly, there are no small features on the remaining volume that necessitate exclusion from the lattice domain. The result, shown in purple in Fig. 3b, is the lattice domain.2184(a)(b)Figure 3: Cross-sections showing (a) the Von Mises Stress result from the component analysis and (b) the lattice domain (purple) with the non-lattice regions of the component domain (gray)21853.8 Unit Cell ExplorationThe original study [46] used the self-supporting Face Centered Cubic (S-FCC) [47] lattice because it is self-supporting and it is easy to model with traditional CAD tools, which allows integration with solid sections of the component domain for analysis. Advances in software have simplified implementation of more complex lattice unit cells, including surface-based lattice structures [14] that can offer improved performance when compared to the S-FCC lattice [47].

This software advancement opens the design space significantly, allowing for selection of whichever unit cell best meets the needs of the specific application.Based on the lattice level requirements and the lattice domain definition, we can define the range of properties that the lattice must achieve. From this, candidate unit cells can be identified and acceptable ranges of lattice parameters for each can be identified. For a uniform lattice, only the maximum value of the lattice level requirements within the lattice volume needs to be used to select a unit cell and its parameters.The Lattice Unit-cell Characterization Interface for Engineering (LUCIE) [28] uses com- piled published data from multiple sources offers a tool for the comparison and selection of lattices. The amount and variety of data needed to make informed decisions for the entire Lattice Vee framework is not available from current literature and must therefore be developed.

The amount and variety of data needed to make informed decisions for the entire Lattice Vee framework is not available from current literature and must therefore be developed. Although the data set currently utilized by LUCIE is limited, it can be used to identify lattices that offer high stiffness and yield stress as a function of the relative percentage of the lattice that is solid, referred to as Volume Fraction (see Fig. 4). For this application, we use a lattice structure derived from the Gyroid Triply Periodic Minimal Surface (TPMS) surface, “thickening” it by adding material equally on both sides of the surface. We denote this lattice as the Triply Periodic Surface (TPSf) of the Gyroid (Fig. 5a); this nomenclature using “f” is used to differentiate the lattice from the zero thickness minimal surface that it visually resembles.

5a); this nomenclature using “f” is used to differentiate the lattice from the zero thickness minimal surface that it visually resembles. This structure was chosen based on the lattice level requirements, the properties available in LUCIE, and manufacturability using LPBF.In Section 3.9, we define the unit cell size and position based on the component geometry, leaving only VF to be defined. This allows us to apply linear elastic homogenization techniques across a range VFs for the Gyroid TPSf to yield a data set that is sufficient to map between material properties and the VF of the lattice to allow for demonstration of this framework [22]. More detail on the homogenization is discussed in Section 3.11.3.9 Lattice Cell DefinitionAs discussed in Section 2.9, the Lattice Vee does not prescribe a single method for the definition of the cells of the lattice.

This allows for the integration of different design and optimization approaches to be used.To fully define the uniform cubic Cartesian TPSf Gyroid lattice selected in Section 2.9, the cell size, orientation, position, and VF must be defined. For this example, we define the lattice density by optimizing the homogenized properties of the lattice. Because the homogenization approach is independent of the component geometry and uses linear FEA, the cell size, orienta- tion, and position are not defined by homogenization and instead defined separately based on the2186Figure 4: LUCIE tool comparing the yield stress of six different lattice structures as a function ofvolume fractioncomponent geometry. The lattice is oriented such that the z-axis of the Cartesian lattice aligns with the cylindrical axis of the component and is positioned such that the cells start at one end of the cylinder.

The lattice is oriented such that the z-axis of the Cartesian lattice aligns with the cylindrical axis of the component and is positioned such that the cells start at one end of the cylinder. The cell size was defined as smallest cell with an edge length greater than 20 mm such that there are an integer number of cells along the cylindrical axis of the component. With the component length of 127 mm, the cell size is 21.16 mm. This size was chosen to give a feasible thickness to lattice features for verification with FEA and production with LPBF. For the optimization, the target of the optimization is based on the component requirements (2.3% of the bulk material). A gradient based optimization scheme was used. A detailed description of the optimization is irrelevant to the primary focus of this work and is therefore omitted.

A gradient based optimization scheme was used. A detailed description of the optimization is irrelevant to the primary focus of this work and is therefore omitted. A VF of 12.8% is needed to meet the component requirements.CxSy +CySz +CzSx = tEquation (1) is the approximation of the gyroid surface used to generate gyroid based lattices.

A VF of 12.8% is needed to meet the component requirements.CxSy +CySz +CzSx = tEquation (1) is the approximation of the gyroid surface used to generate gyroid based lattices. Where x, y, and z are the three axes of a Cartesian coordinate system, t is the level set value used to control the VF of the lattice, Si and Ci are short hand as defined in Equation (2) and Equation (3) respectively where Pi is the periodicity of the lattice in direction i.Si = sin(cid:18)2π Pi(cid:19) i2187(1)(2)(a)(b)(c)Figure 5: (a) Single unit cell of the TPSf Gyroid at 12.8% VF, (b) view from below cell with overhangs that exceed limit highlighted in red, and (c) the normalized Young’s modulus surface for the latticeCi = cos(cid:18)2π Pi(cid:19) iTo generate the TPSf, two level sets of equal magnitude and opposite sign are used.

To achieve the VF target of 12.8%, level set values of t = ±0.199 are used and the lattice was modeled in nTopology [14].3.10 Lattice Cell ManufacturabilityAlthough the manufacturability of the TPSf Gyroid lattice at comparable cell size, VF, and aspect ratio has been demonstrated previously [48], we include an evaluation here for completeness.Only one test of the manufacturability of the lattice must be carried out, because the cells and material distribution are uniform. We show the analysis of overhanging angles relative to an expected build orientation, with the cylindrical axis aligned vertically in the build direction, in Fig. 5b. Area’s that exceed the limit of 40° from the build plate are shown in red. Because the red areas are supported on multiple sides, we consider them as bridges where the bridge lengths are below the limit of 5 mm, per the AM process limitations (see Section 2.2), and are acceptable to print using the LPBF process.

To check minimum feature size, we use the Shape Diameter Function [49] implemented in MeshLab [50] to identify the location of thin regions within the cell, and the wall thickness measurement tool in Netfabb [51] was used to find the minimum thickness to be 0.774 mm. The minimum thickness in the cell is greater than the permissible minimum feature size of 0.5 mm. Also considered here but not explicitly shown are checks for trapped powder, islands, and pinched off features; no concerns were found. The cell thus has passed the manufacturability checks, allowing the process to continue.2188(3)00.10.20.30.40.50.60.70.80.91VolumeFraction00.10.20.30.40.50.60.70.80.91Normalized Property ValueC11C12C44Figure 6: Relationship between homogenized properties of the TPSf Gyroid lattice and volumefraction3.11 Lattice Cell VerificationWe homogenize the lattice as a linear elastic anisotropic material using an in-house code based on methods of Tan et al. [9].

[9]. This homogenization routine provides relationships between the normal- ized components of the stiffness matrix and the VF of the lattice, and results for the TPSf Gyroid lattice can be seen in Fig. 6. The properties of the homogenized material can be compared to the lattice level requirements that fall within the lattice domain, as outlined in the lattice verification plan (see Section 3.6). The normalized Young’s modulus surface of the homogenized lattice cell is shown in Fig. 5c. As the lattice is uniform, from the lattice level requirements and lattice domain, we have the same requirements for each cell of the lattice, namely, that it should be 2.3% of the bulk material response.3.12 Lattice IntegrationThe TPSf Gyroid lattice was modeled monolithically in nTopology Platform, requiring no integra- tion of the cells with one another.

As the lattice is uniformly distributed within the lattice domain, there are no transitions between cells that could produce discontinuities requiring smoothing.3.13 Component IntegrationThe lattice structure can now be united with the solid regions of the component domain (see Fig. 7a). At the intersections of the lattice and solid, we apply a 2 mm radius to reduce poten-2189tial stress concentrations at the transition. Future work should investigate stress cracking at the interface between the solid regions and the lattice structure.This design now undergoes V&V as specified in the planning stage (see Section 3.3), the first step of which is applying FEA on the completed model with the lattice incorporated. Depending on the model complexity and required fidelity, this evaluation can be performed using either a mesh of the lattice itself or by replacing it with a homogenized solid with properties based on Section 3.11. In this application, we mesh the lattice, and Fig.

In this application, we mesh the lattice, and Fig. 7b shows the von Mises stress in the component. Using the yield stress from Table 1, we find that the minimum Factor of Safety is 3.3 which exceeds the required Factor of Safety of 1.5. This design represents a 49.3% reduction in mass compared to the original, non-lattice design.(b) Figure 7: Cross-sectional views of (a) proposed design with lattice structure integrated, (b) resulting von Mises stress from FEA (in Pascals), and (c) the proposed model to be printed(a)(c)With the design successfully passing the FEA validation, we can now generate a model for manufacturing. Our model must also consider the manufacturability of the solid regions, ensuring that they are self-supporting or have sufficient support from either the lattice or from additional support structures. In this example, there are no such issues.

In this example, there are no such issues. To prepare the model, we apply machining offsets to the exterior cylindrical surface, to the four fluid ports, and to the two end faces of the cylinder. We add a larger offset (5 mm) between the component and the build plate to allow for component removal using wire electrical discharge machining. In addition, we add a radius at the interface with the build plate to reduce potential for crack formation caused by residual stresses produced during the LPBF process [52]. We add 45◦ chamfers on the horizontal overhangs of the fluid ports to eliminate the need for support material on those features. Figure 7c shows this2190proposed build plan. Prior to manufacturing, simulation of the manufacturing process can be used to predict deformation and stresses occurring during the process [53, 54, 55]. This can be used to quantify the deformation and determine if it is acceptable (i.e.

This can be used to quantify the deformation and determine if it is acceptable (i.e. that machining allowances are sufficient to account for the distortion and that recoater interference is not predicted [55]).After manufacturing, NDE and pressure testing should be performed to allow for finaliza- tion of the design, that will be integrated into the upper level assembly from which the component requirements were derived.4 DISCUSSIONFor the application of the model in Section 3, Kantareddy et al. [46] did not specify the maximum stress, (or minimum FOS) in their finalized design that was produced, simply that it was below the allowable stress. We can compare the weight savings of the two designs, to show that the design developed here has 6.9% greater savings. This, however, is a function of the factor of safety which was just over twice the limit; further savings can still be achieved.

This, however, is a function of the factor of safety which was just over twice the limit; further savings can still be achieved. Allowing the framework to determine the solid domain and allowing for non-uniform material distribution (as well as other properties of the lattice structure) within the lattice domain would offer the potential for better material utilization, further reducing weight and (possibly) cost.To effectively apply the Vee model, a large database of known lattice structure properties covering impacts of cell size, VF (or other definition of material content), and material is needed. This database would directly inform the Unit Cell Exploration and Lattice Cell Definition steps of the Vee Sections 2.8 and 2.9. This could further include cell aspect ratio, cell shape, and material grading within the cell.

This could further include cell aspect ratio, cell shape, and material grading within the cell. The generation of this database would require a vast design of experiments (DOE) to be conducted either using FEA or a comprehensive experimental campaign for each unit cell considered. The scale of this data collection motivates an open-sourced and collaborative effort to produce consistent data across different efforts.As additional lattice data becomes available in different materials, the requirements for the material to be known prior to starting the Lattice Vee is likely to change, allowing for the unit cell exploration and lattice cell definition steps to expand to encompass material selection driven by the lattice level requirements.5 CONCLUSION AND FUTURE WORKLattice structures offer many benefits for AM including reducing the weight of components. How- ever, the design of lattices is complex and typically considered in a bubble, separate from broader engineering design.

How- ever, the design of lattices is complex and typically considered in a bubble, separate from broader engineering design. We present the Lattice Vee Framework that we propose to close this gap. The framework takes in the component requirements and uses them to inform unit cell selection through requirement flow down that parallels the System Vee. The Lattice Vee offers the ability to bring the design of lattice structures into requirement-driven engineering without limiting the design freedom that lattices afford and enabling wider adoption in engineering design. We then demonstrate the steps of the Lattice Vee through the design, integration, and testing of a component2191with a uniform lattice.

We then demonstrate the steps of the Lattice Vee through the design, integration, and testing of a component2191with a uniform lattice. This example highlights the efficacy of the framework as well the current limitations and future work needed for wide adoption.Further development pertaining to this framework is ongoing and includes integration of a tool chain to produce a streamlined implementation of the Lattice Vee, lattice testing, and development of more quantitative case studies to demonstrate capabilities of the framework that include utilization of different techniques for defining the lattice structure incorporating optimiza- tion and non-uniform lattices. The framework is expected to continue to evolve alongside additive manufacturing design techniques and become more refined through application of the process with case studies and dissemination of engineering data alongside those studies.

We will also explore the application of homogenization, which has been shown to predict the properties of uniform lattices, to non-uniform lattices with different homogenized properties within each cell of the lattice based on the cells unique geometry.To reduce the computational load during manufacturability analysis and homogenization (Sec-tions 2.10 and 2.11), we propose testing a representative set of cells rather then each individual cell of the lattice. For a uniform lattice, all cells are the same and a single cell can be characterized. For a non-uniform lattice, each cell can be unique requiring characterization of each cell individually. For representative cells to be used, clustering of cells must be explored based on similarity metrics. This clustering of cells and selection of representative cells for manufacturability analysis and ho- mogenization has not yet been explored.

This clustering of cells and selection of representative cells for manufacturability analysis and ho- mogenization has not yet been explored. Computational expense, while verifying the performance of lattice cells, could be further reduced by applying machine learning techniques to estimate the properties of lattice cells thereby removing the need to perform homogenization while validating the lattice [56]. This would likely require a training set of data for each lattice topology. This also motivates the data collection effort discussed in Section 4.References[1] Thompson, M. K., Moroni, G., Vaneker, T., Fadel, G., Campbell, R. I., Gibson, I., Bernard, A., Schulz, J., Graf, P., Ahuja, B., et al., 2016. “Design for additive manufacturing: Trends, opportunities, considerations, and constraints”. CIRP annals, 65(2), pp. 737–760.[2] Yang, S., and Zhao, Y. F., 2015.“Additive manufacturing-enabled design theory and methodology: a critical review”.

The International Journal of Advanced Manufacturing Technology, 103(9), pp. 3769–3783.2193[20] Chennamkulam Renjith, S., 2018. “A design framework for additive manufacturing based on the integration of axiomatic design approach, inverse problem-solving and an additive manufacturing database”. Master’s thesis, Iowa State University, Ames, IA.[21] Li, D., Liao, W., Dai, N., Dong, G., Tang, Y., and Xie, Y. M., 2018. “Optimal design and modeling of gyroid-based functionally graded cellular structures for additive manufacturing”. Computer-Aided Design, 104, pp. 87–99.[22] Wang, Y., Hu, D., Wang, H., Zhang, T., and Yan, H., 2020. “Practical design optimization of cellular structures for additive manufacturing”. Engineering Optimization, 52(11), pp. 1887– 1902.[23] Wang, C., Zhu, J., Wu, M., Hou, J., Zhou, H., Meng, L., Li, C., and Zhang, W., 2020. “Multi-scale design and optimization for solid-lattice hybrid structures and their application to aerospace vehicle components”.

M., Gandhi, U. N., Mandapati, R., and Choi, S.-K., 2015. “A design and fabrication framework for periodic lattice-based cellular structures in additive manufac- turing”. Vol. International Design Engineering Technical Conferences and Computers and Information in Engineering Conference.[42] Trudel, E., and ElSayed, M. S., 2021. “Global multiscale design optimization of procedural lattice structures for fatigue enhancement”. Multiscale and Multidisciplinary Modeling, Experiments and Design, pp. 1–23.[43] Osanov, M., and Guest, J. K., 2016. “Topology optimization for architected materials design”.Annual Review of Materials Research, 46, pp. 211–233.[44] Samperi, M. T., 2014. “Development of design guidelines for metal additive manufacturing and process selection”. Master’s thesis, The Pennsylvania State University, University Park, PA.[45] Kranz, J., Herzog, D., and Emmelmann, C., 2015. “Design guidelines for laser additive manufacturing of lightweight structures in TiAl6V4”.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperApplication of the Fog Computing Paradigm to Additive Manufacturing Process Monitoring and Control Muhammad Adnan1,2, Yan Lu1, Al Jones1, Fan Tien Cheng2 1Engineering Lab, National Institute of Standards and Technology Gaithersburg, Maryland 20899, USA 2Institute of Manufacturing Information and Systems, National Cheng Kung University, Tainan, Taiwan, ROC AbstractMonitoring and controlling Additive Manufacturing (AM) processes play a critical role in enabling the production of quality parts. AM processes generate large volumes of structured and unstructured in-situ measurement data. The ability to analyze this volume and variety of data in real-time is necessary for effective closed-loop control and decision-making. Existing control architectures are unable to handle this level of data volume and speed.

Existing control architectures are unable to handle this level of data volume and speed. This paper investigates the functional and computational requirements for real-time closed-loop AM process control. The paper uses those requirements to propose a function architecture for AM process monitoring and control. That architecture leads to a fog-computing solution to address the big data and real-time control challenges.Keywords: Additive Manufacturing, Fog Computing, Monitoring & Control, FunctionalArchitecture, Control Architecture, Data Analytics.Section-1. IntroductionWithin the next decade, the goal of AM is to become one of the primary manufacturing processes.

IntroductionWithin the next decade, the goal of AM is to become one of the primary manufacturing processes. Requirements for achieving this goal include 1) AM software for high-complexity product design and engineering, 2) AM machines for high-quality and low-cost part fabrication, 3) AM sensors for in-situ monitoring of various AM processes, and 4) controllers that can analyze big data and make optimal decisions in real-time [1]. Vendors have been providing similar capabilities to address these requirements for traditional manufacturing for decades. So what is different about AM?AM is an additive fabrication process. AM processes create parts directly from 3D computer-aided-design (CAD) files. For example, powder bed fusion adds, and then melts, metal or polymer powders in a layer-upon-layer fashion. The melting process varies depending on the technology underlying the heat source used in the process.

The melting process varies depending on the technology underlying the heat source used in the process. Despite of the enormous potential benefits of AM, manufacturers are facing a major problem of the production technology: even under the identical process parameters and machine conditions, the quality of the AM products can vary substantially and often unacceptably [2]. Various factors contribute to this problem, including the variability in feedstock materials, process parameters, and build function execution. To reduce the part quality uncertainty, AM process monitoring and control become critical.Unlike the process monitoring of subtractive processes, AM in-situ monitoring relies on multi-modal sensors that generate a large amount of 1d, 2d and 3d data during fabrication. The data are used to estimate the current states of both AM processes and parts, as well as to predict the final states of the parts.

The data are used to estimate the current states of both AM processes and parts, as well as to predict the final states of the parts. The in-situ data can also be used for part qualification and to improve the design and engineering of future AM products.In addition, AM requires multi-loop feedback control, meaning that the measurement and quality monitoring data are used for process control and decision making at multiple sampling rates. AM in-situ monitoring data, generated in real-time from various sensors at different254intervals, becomes feedback to machine controllers. The controllers analyze the data, modify process parameters in real-time or near real-time, or stop the build as necessary.In this paper, we focus on defining a reference architecture for AM process monitoring and control, including 1) identifying the relevant analytics and control functions, 2) proposing a control architecture, and 3) using fog computing to implement them.

In Section 2, we describe the limitation of the traditional hierarchical control architecture. In Section 3, we present a multi-loop control framework for AM process monitoring and control; Section 4 describes the computing requirements for near real-time process monitoring and control functions; Section 5 presents a fog computing based architecture for AM process monitoring and control2. Is ISA 95 the Right, Control Architecture for AMThe ISA ModelFor decades, the ISA 95 model defined by the International Society of Automation (ISA) has been the hierarchical-control choice for traditional manufacturing systems [3]. ISA 95 divides the functions of those systems into levels based on three aligned decompositions: spatial, functional, and temporal. In Figure 1, the triangle represents the ISA functional hierarchical decomposition, with the temporal aspects (the timeframes) of the hierarchy shown on the left.

In Figure 1, the triangle represents the ISA functional hierarchical decomposition, with the temporal aspects (the timeframes) of the hierarchy shown on the left. It is important to note that those timeframes are holdovers from the time when humans performed those functions. The spatial aspects, which are not shown, go from an individual machine to an entire factory. The shaded boxes on the right of Figure 1 show the various software applications that have been developed to implement the functions in each level. That implementation is constrained by the timeframe assigned to each level.Level 2 Moni toring & superv isingLevel 1 Sensing & manipulatingLevel O Production processFigure 1. The ISA HierarchyThe ISA views this functional hierarchy as a control hierarchy, meaning that each function is associated with a controller. That controller’s job is to use the designated software application to manage the execution of that function.

That controller’s job is to use the designated software application to manage the execution of that function. There are rigid rules about 1) the kinds of inter-level “command-feedback” communications and 2) the information exchanges between the software applications that create those communications. As noted above, those rigid rules were originally developed before the computer revolution, and well before the explosion of commercially available, domain-specific, software applications [3].Since ISA’s standardization, both the use of hierarchical principles and the ISA control hierarchy itself have been adopted, and adapted, in many manufacturing domains. The domain- specific software applications, which began to infiltrate the manufacturing sector in 1990s, were run according to the predetermined ISA timeframes. For example, as shown in Figure 1, ERP would run once a day, MES would run once an hour, and so on.

For example, as shown in Figure 1, ERP would run once a day, MES would run once an hour, and so on. Of course, once the required integration problems were solved, these applications could have been run at any time. But, for a variety of reasons mostly associated with keeping the four aspects in Figure 1 completely aligned, they were not.255Does ISA Work for AMThe beginnings of Industry 4.0, new advances in information technology, and a variety of new AI tools are obviating the need for an AM-centric control architecture. A key point of Industry 4.0 systems is decentralized, data-driven decision-making. The purpose is to make each entity more autonomous, with the capability to communicate directly with any other part of the system [3].Advances in information technology include cloud services, camera sensors, fog computing, and edge computing. Cloud services provide a good option for implementing AM design and AM engineering functions.

Cloud services provide a good option for implementing AM design and AM engineering functions. The economic benefits of using cloud services and computing come from efficient resource allocation. Resource management plays a major part in increasing system performance, thereby enhancing user satisfaction [4]. Nevertheless, cloud services naturally clash with both the new Industry 4.0 design principles and the need for reliable, real-time control of AM machines and processes during the “build” function [5].That need depends on the ability to analyze the large amounts of in-situ measurement data being collected by a variety of cameras. Since there are very few physics-based AM process models, that analysis, and hence control, warrants the use of machine learning tools. Many of these tools exist in the cloud. Hence the in-situ data collected at level 1 must also be sent directly to level 3 and 4 functions for off-line analysis.

Many of these tools exist in the cloud. Hence the in-situ data collected at level 1 must also be sent directly to level 3 and 4 functions for off-line analysis. While the cloud services can certainly be used for understanding AM processes and analyzing the final-part’s quality for inspection purposes, they cannot be used for real-time control. For real-time control, in-situ measurement, analysis, and prediction are required to optimize the process parameters and thereby control the machine and process. These real-time or near real-time functions also demand intensive computation power, which cannot be provided by typical edge nodes, such as PLCs or AM machine on-board computing units.In the remainder of the paper, we first propose a multi-loop feedback control architecture for the “build” monitoring and control. We then propose the use of both fog and edge computing as a way of implementing the corresponding functions.3.

We then propose the use of both fog and edge computing as a way of implementing the corresponding functions.3. Multi-Loop AM Process Monitoring and Control FunctionsThe proposed multi-loop AM process and control architecture is shown in Figure 2. It has three major loops: 1) sub-second real-time control 2) Layer wise scan optimization and 3) offline build planning and data driven modeling. Process Monitoring provides multirate data curation functions, including sensing, data acquisition, data fusion and data analytics.The outmost loop “Build Planning” is an engineering task, usually done with human in the loop. The main outputs of Build Planning include 1) a build plan - the process parameters and other information needed to fabricate the product, and 2) the information (models) used to assess the current state of the process and the parts, and to predict the final state of those parts.

The planning process involves two major types of tasks: 1) process specific tasks including lattice, support design and build orientation selection; 2) machine specific tasks to convert the build plan into the formats required by a particular AM machine, usually done by using machine specific software tools.The innermost loop is for real-time control. Based on in-situ feedback measurements, real- time controllers change the process parameters to stabilize process. The feedback measurements can include melt pool temperatures and sizes, sampled at many KHz. And the process parameters include energy power, material feeding rate etc.

The feedback measurements can include melt pool temperatures and sizes, sampled at many KHz. And the process parameters include energy power, material feeding rate etc. As of today, only a few AM machine models have the real-time feedback control capabilities.256build plan lscan path .....AM ProcessProcess Monitoringprocess parameter..Rea~ e Cqntrolone or multiple sampleper-lay ~r ~,:ise -Control nd Planning✓modelper/ayermodelper buildBuild Planningper buildFigure.2 Multi-loop Control Architecture for the Build FunctionThe middle loop “Layer-wise Control and Planning” makes adjustments of scan path and process parameters for the next layer based on the monitoring information from the previous layers.

For PBF, this task has to be done before the recoating process finishes to avoid build time increasing.The multi-loop monitoring and control functions use a variety of machine learning techniques to determine the current states of the process and the parts, to make a prediction about the final state of the part and to generate real-time or near real-time control to bring back the part quality from deviations. The models used for real-time and near real-time control can be learned from past builds and simulations which are usually trained in clouds. The process monitoring functions supporting real-time or near real-time control have to provide process and part state estimation within micro seconds or several seconds to enable the feedback loop. For layer-wise re- planning, multi-modal sensor fusion and model predictive control are involved for defect detection and scan path re-planning.

For layer-wise re- planning, multi-modal sensor fusion and model predictive control are involved for defect detection and scan path re-planning. These tasks are also computational intensive, beyond the capability of the traditional edge computation nodes such as PLCs or embedded controllers. In order to identify appropriate computation infrastructure for the AM process monitoring and control, we will first analyze the functions for in-situ monitoring and control and group them in a function architecture, which will be described in this section.

AM Process Monitoring and Control Function ArchitectureFigure 3 shows a reference function architecture of AM process monitoring and control.Analytics and control functions are grouped and illustrated based on the time criticality.257Off-Line (Minutes Up)3DReconstruction L---=-=.,::..:..:...___JQualhy PredictionNear Real-Time (Millisecond- Second)Melt-pool 8-MterJReal-Time (Microsecond- Millisecond)Anomaly DetectionProcessFigure.3 AM Process Monitoring and Control Function ArchitectureReal-time functions include data acquisition from in-situ monitoring system, measurement data preprocessing - cleaning and tagging, melt pool geometry characterization, fast spectrum analysis of acoustic measurements, anomaly detection, real-time feedback control generation, and emergency reaction, such as build stop.Near real-time functions fuse the process monitoring data obtained from previous layers, conduct layer-wise process and part state evaluation and make decision if the build should be continued or stopped.

If the build shall continue, a scan plan adjustment function may be necessary based on the evaluation or prediction of part quality. With hybrid manufacturing, the layer-wise controller can make a decision to machine off the defected layers and then resume the layer-by- layer process.Offline functions include data analytics for engineering decision. In-situ monitoring and control data from multiple past builds, as well as the part development lifecycle data, are aggregated for analysis. The data can be used for training to correlate process settings with process signatures, microstructure properties and mechanical properties. The resulted models are used for build plan generation. Machine learning can also be used to train models which are used in real or near real-time process monitoring and control.In this paper, we focus on the analysis and prediction functions for in-process control.

Many of the tasks associated with these functions are shown in Figure 3, including tasks in the real-time group and tasks in the near real-time-group. In the remainder of this paper, we will discuss the near real-time functions, which demand the use of fog computing for implementation. Near Real-Time FunctionsThe main purpose of the near real-time process monitoring and control functions is for layerwise decision making. Multiple commercial AM systems and third party in-situ monitoring258systems are equipped with such capability. The cycle time for layerwise decision making varies from seconds to minutes, depending on the design of the parts, the scan patterns and the process parameters. The middle layer of Figure 3 lists the mostly reported layerwise data analytics and control functions.

The middle layer of Figure 3 lists the mostly reported layerwise data analytics and control functions. They are sensor data fusion, residual stress estimation, 3d model reconstruction, defect detection, process and part quality prediction, and layer parameters optimization.Multiple in-situ sensors are being utilized in AM machine for monitoring and controlling AM part quality. The sampling rates and volumes of the sensor data vary. For real-time monitoring and control, these sensors data such as temperature, melt pool geometry, acoustic signal, process parameters, etc. need to be registered correctly and fused effectively. Sensors data fusion is a major area of research for better understanding of the AM process [6, 7].3D in-situ measurement such as Optical coherence tomography (OCT) is used for both surface and inner-structure defect detection, as well as dimension analysis.

Due to its high resolution and non-destructive nature, it also useful for surface-void detection, loose powder detection, and subsurface feature detection. OCT is also used to find cracks and un-melted powder areas during production [8, 9].Accurate residual stress estimation is a key step in attaining maximum dimensional accuracy and avoiding early fatigue failure [10]. Many process parameters affect the residual stress in AM [11]. The estimated residual stress based on in-situ thermal measurements can be used as the inputs to re- plan scan paths and optimize process parameters for the next build layer. Therefore, they work as a promising tool for AM part quality control.Finding defects using layerwise images provides timely control over the ongoing manufacturing process. By detecting a catastrophic defect during the process, we can stop a failed build early, reduce the cost and avoid the waste.

By detecting a catastrophic defect during the process, we can stop a failed build early, reduce the cost and avoid the waste. There are many existing methods to find defects such as ANN, Bayesian classifier, support vector machines (SVM), and Convolutional Neural Networks (CNN). CNNs earned attention due to the accuracy and fast execution time as compared to other methods. The execution speed of a CNN is enhanced through the use of high-performance computing resources such as advanced GPU. The output of the function can be used for (layer- wise) process control, supplementary process decisions, or remedial actions [12-14]. Some researchers use an acoustic signal for in-situ quality prediction in AM using deep learning [15-16]. Automatic virtual metrology (AVM) is a technique in which the quality of the manufactured part can be predicted without actual manufacturing by utilizing previously obtained in-process measurements.

Automatic virtual metrology (AVM) is a technique in which the quality of the manufactured part can be predicted without actual manufacturing by utilizing previously obtained in-process measurements. For AM quality prediction, the AVM system is based on in-situ data sensed during prior manufacturing [17-18].Layerwise parameter-optimization functions re-plan the process settings for the subsequent layer. These functions consider outputs from the quality prediction, OCT, the residual stress estimation, and the defect detection functions, and perform path re-planning using ANNs or genetic algorithms [19].Today’s AM machines are instrumented with a myriad of sensors. In addition, many of 3rd party provided sensors can be easily integrated which generate and produce data over extremely short periods of time.

In addition, many of 3rd party provided sensors can be easily integrated which generate and produce data over extremely short periods of time. Not like the traditional machine tools using 1d sensors, many of today’s AM-sensors are generating vast amounts of 2d (image data) and even 3d data at increasingly faster rates – much different than the traditional numerical data. Tens of thousands images from high speed co-axial cameras are generated each layer. In addition, staring cameras produces multiple high-definition global view images each layer. Fusing the images with the machine control commands and other sensor measurements are critical for layer-wise control, planning and259decision making. The computation requirements to perform the data analytics and control are very high.4. Computing Requirements in AMComputational speed is the major issue for the real-time control system.

The computation requirements to perform the data analytics and control are very high.4. Computing Requirements in AMComputational speed is the major issue for the real-time control system. The traditional method can deliver precise melt pool measurement but deficiencies of processing efficiency. For real-time control in AM, a major problem is that current single node systems cannot meet the real- time computing requirements. For example, a traditional melt pool classification method takes 3.5 milliseconds to process a melt pool image. If a co-axial camera generates approximately 3000 images each layer, classifying the full layer of melt pool images would require approximately 10 seconds. This calculated time does not include the data transfer time; sending such a large amount of data to the cloud for processing will require additional time for transferring and processing.

This calculated time does not include the data transfer time; sending such a large amount of data to the cloud for processing will require additional time for transferring and processing. Due to these problems, currently manufacturing automation systems fail to meet the AM requirements for real-time monitoring and control [20].The computing requirements for AM processes are dictated by the volume, velocity, variety, and veracity of the data. According to Wang et al., there are up to 2.3 trillion voxels in a typical build volume; about 600 variables logged during AM processes on a per seconds basis, giving up to 300 MB of data per build; and up to 0.5 TB of data collected per build using in-situ monitoring [21]. Photodiodes and pyrometers are the most common devices used in SLM monitoring systems due to their fast response and easy integration. Both devices boast acquisition rates above 50 kHz.

Photodiodes and pyrometers are the most common devices used in SLM monitoring systems due to their fast response and easy integration. Both devices boast acquisition rates above 50 kHz. If process is monitored with a high-resolution camera (20 kHz or more) for a few seconds, even for very short time interval, several GB of data per second can be collected.Figure 4 shows the correlation between camera frame rate and resulting data. Clearly the high volume of data cannot be transferred using current industrial interface communication protocols [22]. For real-time data acquirement and transfer, the data rate may not exceed theAM Monitoring System Data Rate10~+ ~ - - - - - - - - - - - - - - - - - - -~ - - - .Figure.4 Ratio of FPS to Data Ratelimit defined by communication protocol; for example, USB3.0 can handle a maximum rate of 640 Mbps. Studies show that data rate is fairly higher than range of the current available260communication protocol [22].

Studies show that data rate is fairly higher than range of the current available260communication protocol [22]. AM in-situ monitoring data processing for near real-time control is a big data problem. In the following sections we propose an AM process monitoring and control architecture based on fog computing.5. Fog Computing and ArchitectureOwing to the amount of data and the data’s dynamic nature, processing it within the time constraints required for real-time applications is a big challenge in AM. In the AM process, the huge amount of data produced in a very short time leads to time constraints and latency problems. Both problems make it impractical to use a cloud-computing architecture for real-time, process- control applications.The cloud model can be organized using three major concepts: infrastructure as a service, platform as service, and software as service.

Some of the main benefits of the cloud model are virtualization, scalability, on-demand self-service resource pooling, and location freedom. In contrast with cloud-based manufacturing, fog computing shifts the heavy data workload from the centralized cloud to near edge devices. Due to the latency problem, the amount of data communicated using the cloud must be reduced for online, real-time machine monitoring, diagnosis and control. In fog-empowered architectures, manufacturers can save sensitive data on local machines while using the intelligence, data analysis and training applications provided by high-performance, cloud computing [23]. Fog Based SystemsFog computing works as an intermediate layer between IIoT (Industrial Internet of Things) and cloud for more responsive services. In fog computing, each computer node works independently to provide intelligence on the outer edge of ubiquitous networks, without requiring a persistent network connection.

In fog computing, each computer node works independently to provide intelligence on the outer edge of ubiquitous networks, without requiring a persistent network connection. This reduces network traffic and enhances scalability and security [24-25]. Fog computing is a geographically distributed computing architecture with a resource pool consisting of one or more ubiquitously connected heterogeneous devices (including edge devices) at the edge of the network and not exclusively seamlessly backed by cloud service. The architecture provides elastic computation, storage and communication in isolated environments to several clients in proximity [26].Wu et al. proposed a real-time remote computational framework based on fog architecture for process monitoring and prognosis in cyber physical system (CPS). That framework utilized the wireless sensors, cloud, and machine learning [27-28].

That framework utilized the wireless sensors, cloud, and machine learning [27-28]. Fog was shown to perform well in smart city applications including data demonstration, feature extraction, anomalous and hazardous event classification, and security measures. A fog-based distributed architecture was proposed to support the data collection and analysis in fast response applications [29-30]. A combination of PMML-encoded ML models and fog computing was proposed for gaining the key ideas of decentralization, security, reliability, and privacy in Industry 4.0 [31]. We could not find any literature on fog-based functional and control architectures specialized to AM.Fog-centered computing models are currently utilized more and more to fulfill the requirement of IIoT, CPS, and mobile computing. It is suitable for applications where instant feedback and response are required. To handle time limitations, fog computing architecture plays an important role in real-time application [32].

It is suitable for applications where instant feedback and response are required. To handle time limitations, fog computing architecture plays an important role in real-time application [32]. In the remainder of this paper, we will discuss our proposed architecture to deal with AM, time constraints. Proposed Fog Based ArchitectureThe proposed architecture is a combination of edge, fog, and cloud methods, as shown in Figure 5, which mirrors the temporal decomposition of control functions shown in Figure 3. As shown in Figure 5, The Intelligent In-Situ Monitoring and Computing System (IIMCS) comprises261five major layers: sensors, intelligent edge per sample, intelligent fog per layer and per build decision making, and cloud.

These sensors capture huge amounts of image data (on the order of gigabytes) within an extremely short period of time (on the order of seconds). To 1) transfer this amount data to the cloud, then 2) to analyze the data using some kind of AI tool, and, finally 3) to return it to AM controller who then must make a decision will certainly take more than the microseconds allowed for the analysis of each sample (on the order of microseconds). Consequently, utilizing service to perform that analysis, the cloud is not possible – primarily due to communication delays (latency). The reason is that no currently available communication protocol supports the transfer of that amount of data involved within the time constraints required for real-time AM quality control.As an example, we describe a concept for a new, open-source, smart camera that eliminates the need to transfer image data to the cloud (see Figure 6 left side).

Our smart camera is an edge device that both collects and analyzes melt-pool images in real time. Analysis operations are performed as part of the real-time, monitoring functions shown in Figure 3. The smart camera performs various different kinds of operations, such as anomaly detection, preprocessing and feature extraction, on the images obtained in monitoring the melt pool.If no anomaly is detected, the smart camera outputs the feature data to the fog layer for further analysis. If an anomaly is detected, the relevant raw data is sent to the fog layer for further analysis.

It then uses the data to control multiple systems including the scanning system, the environmental system, and in-situ systems [33]. There are many, third-party, commercial, in-situ control systems available such as SLM solutions, Stratonics , EOSSTATE, QM melt pool 3D and Sigma Labs [34]. To achieve successful real-time control, data from the AM system, as well as from any utilized commercial and custom systems, needs to be fused for further analysis.This fog component deals with the layer-by-layer data of the AM process. The area of the melt pool and correlates that estimate with the current laser power and scan speed. Layerwise data analytics provides the information needed to adjust the parameters before the next fabrication layer begins, if that information indicated any kind of problem in the process. Our design provides feedback in real-time.

Our design provides feedback in real-time. Thus we need to analyze the layerwise data and provide suggestions or warnings to a feedback controller for better process parameters adjustment as soon as possible. Layerwise analysis function data transfer to the parameters adjustment as soon as possible. Layerwise analysis function data transfer to the next layer for build level decision.Intelligent Fog Per Build Decision MakingThe intelligent fog per build consists of knowledge decides whether to continue or stop the process, based on current data. This component controls the main functionalities of the proposed263architecture. The designed architecture performs all these tasks locally within the boundary of the factory, which enhances the trust of manufactures in term of data security. Due to privacy and security concerns, technology manufacturers do not want to send their data outside the factory.

Due to privacy and security concerns, technology manufacturers do not want to send their data outside the factory. Why Fog-based Architecture? The proposed architecture is a combination of cloud and fog services. The advantages of fog computing are location awareness, mobility, low latency and physical dispersion. Fog computing is not an alternative to cloud computing. Rather, it eliminates the drawbacks of cloud computing and improves the efficiency [4]. Fog computing has the following advantages compared to cloud- based solutions: [35, 36]1) Latency ConstraintFog computing has the capability of solving any latency issues by performing data analysis, control and time-restricted tasks near to the end user. Cloud technology often fails in time-restricted applications because it works in a centralized fashion; all processing is done in a distant location requiring a substantial amount of time to transfer data, process data, and send it back to the user’s location.

In AM quality-control processes continuously generate huge amounts of data and require near instantaneous feedback for effective quality control. Consequently, typical Cloud technology, owing to its substantial latency, is not a feasible computing choice for AM control.2) Network Bandwidth ConstraintFog supports hierarchical data processing in conjunction with the cloud. It creates balance among application conditions and available computer resources by allowing different processing at different levels or locations. In this way, fog reduces the quantity of data transferred to the cloud that requires higher bandwidth. The frame rate is very high for in-situ AM measurements, making it impossible to transfer these large amounts of data in a very short period to the cloud. Fog technologies, however, have the capability to perform operations on large amounts of data without transferring them to the cloud.

Fog technologies, however, have the capability to perform operations on large amounts of data without transferring them to the cloud. In the suggested architecture, the fog dedicated layer handles the complex task of data allocation for real-time analysis.3) Resources Constrained DevicesCameras and pyrometers are examples of resource-constrained devices because they possess very small amounts of both processing power and storage memory. That means the data they collect must be communicated somewhere and processed very quickly. These devices typically are incapable of sending data to the cloud due to reasons like power, bandwidth, and cost. Fog address this limitation.4) Uninterrupted Service/Irregular Connectivity with CloudFog systems provide services even when there is an irregular connection withthe cloud.5) Security ChallengesFog systems work as proxies for all limited-resource devices for software and security authorization updates.

Further, as limited-resource devices have poor security capabilities, the fog system can provide antivirus scanners.6) Reduce data movement across the networkBy limiting data movement across the network, fog systems significantly reduce network congestion, remove the drawbacks of centralized computing systems, allow the necessary data to stay closer to the end user, and provide enhanced scalability arising from the use of virtualized systems.2647) Removes the core computing environmentBy removing the central computing setting, fog systems make real-timequality control in AM manufacturing possible.8) Faster responseIn addition to providing a sub-second response to end users, it alsodelivers high levels of scalability, reliability and fault tolerance.ConclusionThe proposed fog-based architecture addresses the challenges of the real-time quality control for additive manufacturing using big data.

The fog architecture shifts the high bandwidth and latency-sensitive processing to near-the-edge devices in order to perform necessary functions without sending data to the cloud. The intermediate layer of fog solves the problems of insufficient bandwidth and latency for real-time data analytics and decision making applications in the AM process. The proposed architecture also helps the machine builders, in-situ system providers, and software vendors to understand the function requirements for better integration of their products and services. In future work, the proposed control architecture will be prototyped in a research environment for validation, and performance will be measured and shared with the AM community.DisclaimerNo approval or endorsement of any commercial product by NIST is intended or implied. Certain commercial equipment, instruments or materials are identified in this report to facilitate better understanding.

Certain commercial equipment, instruments or materials are identified in this report to facilitate better understanding. Such identification does not imply recommendations or endorsement by NIST nor does it imply the materials or equipment identified are necessarily the best available for the purpose.AcknowledgmentThis work was partially supported by the “Intelligent Manufacturing Research Center” (iMRC) from The Featured Areas Research Center Program within the framework of the Higher Education Sprout Project by the Ministry of Education (MOE) in Taiwan. References1)2)3)4)5)6)Vandone, Ambra, Stefano Baraldo, and Anna Valente. "Multisensor data fusion for additive manufacturing process control." IEEE Robotics and Automation Letters 3.4 (2018): 3279-3284. Frazier, William E. "Metal additive manufacturing: a review." Journal of Materials Engineering and Performance 23.6 (2014): 1917-1928. Åkerman, Magnus. Implementing Shop Floor IT for Industry 4.0.

lid Freeform fabrication Symposium Austin, Texas, 2010Application of Topology Optimization in Modern Additive Manufacturing Gilbert Chahine1, Pauline Smith2, Radovan Kovacevic11 Research Center for Advanced Manufacturing, Southern Methodist University, Dallas, Texas 75205 2 The Army Research Laboratory, Aberdeen Proving Ground, MarylandThe current work examines the principle of topology optimization (TOP) and solving the problem of minimal compliance, and its applications in modern additive manufacturing (AM). The theory of TOP has been excessively investigated for the past few decades; however its practical use is hindered by manufacturing limitations of manifesting the designs into physical parts. TOP numerically determines the favorable topology configuration contained in a workspace that is constrained by a specific set of external supports and applied loads.

TOP numerically determines the favorable topology configuration contained in a workspace that is constrained by a specific set of external supports and applied loads. It overcomes conventional design traditions, and solves for finding the path of least resistance, mimicking phenomenon found in natural structures, such as bone and wood. The present research focuses on the different methodologies invested in the field, and also inspects the possibilities and efficiency of using numerical design, which is powered by TOP, in modern AM processes such as Electron Beam Melting enabling the production of highly complex parts. The major interest is the production of functionally graded porosity (FGP) in bio-adaptable dental implants able to reduce stress shielding, promote faster osseointegration, and provide micromotion between the dental prosthesis and the root form of the implant.

Future implications of the field include optimized porous structures aiming towards application-defined stiffness and pore size, enabling a digital design with advances analogous to modern AM.IntroductionTopology optimization (TOP) is a relatively new field of study. It deals with the mathematicalproblem of finding the optimal distribution of material of a component based on specific criteria of performance or physical aspects. Although some of the earliest principles date back to 1904[1], the field have not received any remarkable improvement until the emergence of computational power and its maturity in the mid-1980’s. Analogous to biomatter in nature, TOP is believed to portray a more goal oriented design routine when compared to conventional means where human intuition and experience plays the most crucial role in the preliminary design conceptualization of a structural component.

It is the optimal solution that provides a more efficient use of a material by maximizing the performance of user- set resource parameters such as mass or cost, or by minimizing the needed resources to accomplish a user- set performance.Two major approaches of representing topology prove mostly eminent: a voxel-based topology, and a truss-based topology. The current work will examine both approaches while exhibiting an example where TOP is used as a design tool which applies to the design of lattice structures and their consequent building using an AM process; i.e. electron beam melting (EBM).According to Li et al., TOP deals with identifying the areas of a domain that have inefficient levels of stress , and then removing them in order to increase the overall efficiency. By doing so an optimal material distribution would emerge [2].

By doing so an optimal material distribution would emerge [2]. Li elaborates two approaches to achieve the aforementioned goal:Solid Freeform fabrication Symposium Austin, Texas, 2010the homogenization method [3], and the discrete variable method which he credits to the work of Shim and Manoochehri in 1997 and Xie and Steven [2]. The former requires complex and cumbersome mathematical techniques, and proving the existence of a solution was the main objective of the corresponding studies instead of practically finding the solution. The latter approach, which employs discrete methods, takes advantage of computational techniques where an evolutionary scheme is used in parallel with a finite element model to find the optimal distribution of material.

The next few sections will portray and discuss the mathematical model of one of the most common TOP objectives (minimal compliance problem), and its corresponding implementation into a Matlab code by Sigmund et al [4], the expansion of the Matlab code into solving three dimensional problems, using ANSYS® to solve TOP problems to produce an optimal design for a lattice structure under uniaxial compressive stress, and finally implementing truss-based TOP to produce an optimal lattice structure designed to perform under a tri-axial strain field.Minimal Compliance problemThe minimal compliance problem, also known as the maximum global stiffness problem, isperhaps the most common TOP approach.

A domain in either R2 or R3 governed by a set of constraints, loads, and a user-defined volume reduction is evaluated for the optimal material distribution (See Figure 1).Figure 1: Schematic of a topology optimization environmentAccording to Sigmund and Bendsøe, from a discrete model point of view where the domain is meshed into discrete elements, the mathematical formulation of the problem at hand is as follows [5]:(1)where c is the compliance tensor, f the force field, and u the displacement field. The minimization of the compliance is governed by a common finite element method relationships used in solid mechanics which includes the force and displacement fields, as well as the stiffness matrix which is affected by the change of density. The solution of Equation (1), results in a distribution of densities throughout the domain, which after several iterative steps will converge towards the minimal compliance.

The solution of Equation (1), results in a distribution of densities throughout the domain, which after several iterative steps will converge towards the minimal compliance. However, a binary solution is strongly desired, where a density of 1 would indicate the presence of material at a certain element, in contrast to a density of 0 or, in avoidance of singularities, a minimal density value, would indicate the presence of a void at the corresponding element. Consequently the binary solution outlays the optimal topology of a structure.Solid Freeform fabrication Symposium Austin, Texas, 2010In order to enforce a binary solution, a solid isotropic material with penalization model (SIMP) has been suggested; where a penalization factor is added to relay the relationship between a material’s density and its stiffness, in order to lead towards a binary solution of the minimal compliance problem [6, 7, 8].

Equation (2) portrays the mathematical formulation of SIMP [5]:( )( ) ( )(2):∫ ( )( )where ( ) corresponds to the stiffness at a certain indexed element, ( ) is the density of the element, has been proven to be physically permissible corresponding to a minimal value of penalization which depends on the material’s Poisson’s ratio [9]. By implementing the SIMP on the minimization equation of Equation (1), the modified objective function to minimize becomes Equation (3):( ) is the stiffness of the base material, and p is the penalization factor.

This relationship(3):c =where x, corresponds to the density field, which after convergence corresponds to the solution of the mathematical problem.Three main approaches for solving the SIMP minimization problem can be citedThe method of moving asymptotes (MMA) [10]  Sequential linear programming (SLP)  The method of optimality criteria (OC) [11]Sigmund presented a paper in 2001 [4], where a 99 line Matlab code is used to solve two-dimensional static TOP problems, with the options of using multiple load scenarios. The code employs the SIMP model, as well as the OC method for solving (See Figure 2 ). In addition it uses a mesh independency filter to prevent check board areas in the solution.Figure 2: OC method employed by Sigmund [4]Figure 3 and Figure 4 displays three examples of solutions found using Sigmund’s code. The former portrays a cantilever beam with a load at the middle.

The former portrays a cantilever beam with a load at the middle. By using a symmetry approach the number of elements can be reduced, hence improving computational efficiency. As portrayed in the figure,Solid Freeform fabrication Symposium Austin, Texas, 2010computing time was 25 seconds and the result is a familiar topology found in the design of suspended bridges.Figure 3: An example of a topology optimization using the code provided by Sigmund et al. [4], and a modern suspended bridge (courtesy of ©Trent Parke/Magnum Photo)Solutions generated by TOP depend on different factors such as mesh resolution, methods used for solving, and iterative/evolutionary parameters. Effects of such factors include checkboards and mesh dependent designs. Some literature surveys the occurrence of such problems and the available techniques to resolve them [12].

Effects of such factors include checkboards and mesh dependent designs. Some literature surveys the occurrence of such problems and the available techniques to resolve them [12]. Bourdin presents an example of a filter used to deal with these problems to ensure a sharp and valid solution [13].Figure 4: Additional Examples of a topology optimization using the code provided by Sigmund et al [4]An extension to conventional TOP which deals with static loads, is dynamic TOP, where instead of just maximizing compliance, it becomes of interest to maximize fundamental eigenvalues, where resonance of the structure is to be avoided by making sure that its natural frequency is higher than any loading frequency imposed any kind of dynamic loading device.

After examining the code and methods employed by Sigmund, the current paper portrays the implementation of a Matlab code and algorithm analogous to the former in order to solve three-dimensional TOP problems.Implementation of a voxel-based three-dimensional TOP algorithmThe first step for implementing the three-dimensional analysis is to define the unit element that isused to mesh the domain of interest. A structured brick element similar to elements used in commercial finite element software is established, constituting of the 8 nodes and 24 DOFs (See Figure 5).Solid Freeform fabrication Symposium Austin, Texas, 2010Figure 5: Schematic of the 8 node, 24 DOF, brick element used in the algorithmThe next step is to generate the stiffness matrix associated with the brick element. After examining the existing literature, no resources providing the matrix are found.

After examining the existing literature, no resources providing the matrix are found. However, employing principles of modern solid mechanics [14], and using Mathematica® to perform some parametric calculations the matrix is generated following the set of Equations in (4).(4):where [E] is the 6x6 stiffness matrix of the material. [B] is a 6x24 matrix found by the differentiation of [N] which depends on the geometrical aspects of the element. The result of the generation is [k] the 24x24 stiffness matrix of an 8 node brick element. The algorithm used for the three-dimensional TOP is analogous to the one used in Sigmund’s code. A TOP function solves the minimal compliance problem employing an OC method.

The algorithm used for the three-dimensional TOP is analogous to the one used in Sigmund’s code. A TOP function solves the minimal compliance problem employing an OC method. The condition for optimization, which follows an FEM relationship, is implemented using the [k] matrix generated earlier.Solid Freeform fabrication Symposium Austin, Texas, 2010Figure 6: Algorithm used in the three-dimensional topology optimizationThe user inputs the dimensions of the domain, a penalization factor value, and a volume fraction to the algorithm, along with loading conditions and constraints. The result is a binary distribution of material, where a “minimal value” corresponds to void, and a maximum value of “1” corresponds to material.

The result is a binary distribution of material, where a “minimal value” corresponds to void, and a maximum value of “1” corresponds to material. Figure 7 and Figure 8 display two examples of TOP where a domain containing 1000 brick elements is loaded under an axial and shear force respectively.Figure 7: Topology Optimization of a beam under a shear forceThe time of convergence is in the order 103 seconds, and it requires over 20 iterations for the solution to converge. Increasing the definition of the mesh, producing more elements, leads to a severe increase of computing time, hindering the practicality of employing the developed code to produce three dimensional topology with decent resolution.

Nonetheless, the code proved very helpful to explore an application aspect of TOP, similar to Sigmund’s code which as mentioned before deals with two-dimensional topologies [4].Solid Freeform fabrication Symposium Austin, Texas, 2010Figure 8: Topology Optimization of a beam under axial forceTopological Optimization using ANSYS®Due to the educational aspect of the code produced by Sigmund, and the consequent three- dimensional implementation, the code lacks the required computational efficiency to handle higher resolution and a larger number of elements. In order to produce topologies with decent resolution that can be built using AM techniques, a commercial TOP software is used in the current work; i.e. the topological optimization module in ANSYS®. Figure 9 displays the model used for the TOP. The main objective is to produce a structure that can be periodically replicated in order to span over a volume.

Figure 9 displays the model used for the TOP. The main objective is to produce a structure that can be periodically replicated in order to span over a volume. In order to do so, the resulting topology has to exhibit symmetry in the 3 axial directions. The mentioned model represents 1/8th of the total volume of the topology. Material models are established with a Young’s Modulus of 110 GPa, and a Poisson’s ratio of 0.33 which corresponds to the mechanical properties of Ti6Al4V. The cube is meshed into 24 389 SOLID95 10 node elements. A zero-displacement constraint is imposed at one of the corners of the cube, while a force distribution in the Z direction is imposed to the diagonally opposing corner.Figure 9: Model, mesh, loads, and constraints for the domain used for the ANSYS® topological optimizationThe model is solved for a static structural solution, to make sure that all the parameters in the model are properly defined.

A topological optimization is executed, with a converging error <0.001, using the OCSolid Freeform fabrication Symposium Austin, Texas, 2010method to find the optimal topology for a 50% volume reduction. The solution takes 4425 seconds to converge, after 4 iterations.Figure 10: Density plot after topological optimization (left) and the isolated elements with density over 0.4 (right)Figure 10 displays the resulting density plot of the optimization. By identifying the nodes with a density exceeding 0.4, and the attached elements, the optimal topology of the structure emerges.

By identifying the nodes with a density exceeding 0.4, and the attached elements, the optimal topology of the structure emerges. The next step is to export the nodes into Geomagic©; a scanning data processing software which specializes in reverse engineering and pre-CAD preparation.Figure 11: Extrapolation of the result of TOP which represents 1/8th of the domain into the full domain constituting the unit cell of a lattice structureAfter a series of volume wrapping, surface smoothing, and triangle count reduction operations, the 1/8th of the unit cell is replicated and joint to form a unit lattice cell, with either an open cell or closed cell configuration. The resulting cell is a voxel-based topology optimized unit, which can be extrapolated to produce porous material, and lattice structures that exhibits maximum stiffness in the uniaxial direction ofSolid Freeform fabrication Symposium Austin, Texas, 2010the predefined load.

By changing the volume fraction reduction, altering the level of porosity, the optimal stiffness of the structure will vary, hence providing the user control over the mechanical performance of the material.Truss-based TOP An alternative approach to the aforementioned voxel-based TOP is a truss-based TOP. The basic method for solving the problem is referred to as the ground structure method [5]. The technique deals with setting a fixed number of nodes, and then determining every possible truss connecting the nodes to each other. The areas of the trusses are then varied according to the loads and constraints, with a possibility of the areas being zero, or a minimal value to prevent singularities. A minimal area means that the truss is unnecessary and is consequently removed. Figure 12 display the possible 2016 trusses according to a 4*4*4 node matrix.Figure 12: All possible 2016 connections of a 4x4x4 unit cellWork done by Sigmund et al.

Figure 12 display the possible 2016 trusses according to a 4*4*4 node matrix.Figure 12: All possible 2016 connections of a 4x4x4 unit cellWork done by Sigmund et al. in 1995, deals with finding the optimal truss topologies according to a strain field = [1 1 1 0 0 0] [15]. The truss topologies displayed in Figure 13 are analogous to the one found by Sigmund; however, keeping the cross-sectional areas constant throughout the unit cell for simplicity. The topologies are designed using netfabb, software that specialized in the design of lattice structures and their extrapolation into solid designs for AM preparation. The four cell structures are different, even though they follow the same strain field. However, some added criteria drove towards the various designs. For example cell A has no trusses on the outer surface of the domain cube. In contrast cell B only has truss connecting nodes on the outer surface.

For example cell A has no trusses on the outer surface of the domain cube. In contrast cell B only has truss connecting nodes on the outer surface. Cell A and cell D have an isotropic geometrical distribution, while cell C and cell B manifest anisotropy. In addition, cell D is a merger of three cells of type C oriented in three different orthogonal directions.Solid Freeform fabrication Symposium Austin, Texas, 2010Figure 13: 4 different optimal truss topology solutions corresponding to a strain field e= [1 1 1 0 0 0], found in Sigmund’s work in reverse homogenizationLattice Structures Designed by TOP by EBMOne of the main objectives of the current work is to find the application of TOP in AM. Due tothe abilities of modern technologies of rapid manufacturing to produce parts with highly complex geometry, the use of TOP to produce optimal geometries according to a wanted performance proves to be intuitive.

In the case of the design and production of lattice structures, the few previous steps displayed in this paper can be used as design inputs to produce the corresponding lattice structures (See Figure 14).Figure 14: TOP, a design input to lattice structureFunctionally graded porosity (FGP), a concept introduced by previous work [16] deals with locally tailoring the microstructure of a material in order to convey a locally tailored mechanical performance.

FGP is one of the major design features found in the concept of bio-adaptable dental implants, where the variant mechanical performance profile plays an important role of matching the mechanical properties of the surrounding bone; hence, preventing stress shielding and promoting healthier osseointegration.Solid Freeform fabrication Symposium Austin, Texas, 2010Figure 15: Process track to implementing TOP for FGP designLattice structures produced by EBM following TOP design can be implemented to FGP following the process depicted in Figure 15. The lattice structures designed using TOP, are built using EBM and then mechanically tested for validation. Design variables such as truss aspect ratio, or porosity are altered in order to see their effect on mechanical performance. The results of the mechanical testing are then statistically tabulated and analysed.

The results of the mechanical testing are then statistically tabulated and analysed. The tabulated data can then be used to design FGP, where a user defined mechanical performance distribution is matched with the corresponding geometrical features. For example, in the case of bio-adaptable dental implants, one design scenario suggests that the surface stiffness profile of the implant matches that of the surrounding bone, which similarly to any biomatter has anisotropic mechanical properties. The stiffness profile is detected using micro-computed tomography (µCT) scans, where the gray scale values of the images are directly dependent on bone’s density and stiffness.Figure 16: Three different optimal structural designs, extrapolated into 20 mm lattice cube, and built using an EBMThe cell types A, B, and C of Figure 13 are scaled to 2 mm unit cells and then extrapolated to cover a volume of 800 mm3 using the netfabb software.

The designs are then built using and ARCAM EBMSolid Freeform fabrication Symposium Austin, Texas, 2010A2 machine, using 300 mesh size Ti6Al4V powder, with a layer thickness of 100 µm (See Figure 16). The corresponding cubes will go through mechanical testing to characterize the mechanical performance of the three different cell types. The biggest pore size in type C cells measured 1.42 mm, as displayed in Figure 17, which is slightly smaller than the design pore size of 1.6mm.Figure 17: Micoview® snapshot of the lattice structure with axial front lighting (left), and back lighting (right)SummaryThe current work presents a brief overview of TOP that covers voxel-based and truss-based optimization and their potential application towards the production of lattice structures and FGP.

Contemplating TOP as a numerical design tool completes the loop of digitizing the modern track of design and production of components where the potential of material performance reaches a more efficient position. TOP enables the design of features with highly complex geometrical aspects of porosities, where AM enables the production of such complex parts. These features play an essential role in improving the performance of biomedical implants, where the customization of implants matches the singularities found in every patient’s case, contributing toward longer implantation life and the improvement of patient satisfaction and well-being.Acknowledgement The US Army Research Office granted SMU a grant to develop a program entitled: “Advanced Precision Supply Parts Manufacturing”, Grant No. W911NF-07-2-0056.

W911NF-07-2-0056. In October 2007, in corporation with the Army Research Laboratory, SMU acquired the ARCAM A2 machine, for solid freeform fabrication by electron beam in a vacuum environment manufactured by ARCAM AB, Gothenburg, Sweden.The authors would like to thank Professor Ole Sigmund for his kind assistance concerning some inquiries regarding his work and Magnum Photos for providing the photo in Figure 3.Solid Freeform fabrication Symposium Austin, Texas, 2010References[1]. Rozvany, G.I.N, “Aims, scope, methods, history and unified terminology of computer-aided topology instructural mechanics.”, Struct Multidisc Optim 21, 90-108,Springer-Verlag 2001[2]. Li, Q.; Steven, G.P.; Xie, Y. M. 1999: On equivalence between stress criterion and stiffness criterion inevolutionary structural optimization. Structural Optimization 18, 67–73[3]. Bendsøe, M.P.; Kikuchi, N. 1988: Generating optimal topologies in optimal design using a homogenizationmethod. Comp. Meth. Appl.

PLICATIONS OFLASER DIGITIZING AND SURFACING TECHNOLOGIES"reverse engineer" an already existing part for which there is no existing CAD model, print,Many companies have manufacturing requirementsto copy orpattern, or mold. In the past this process has been very tedious requiring weeks and sometimes months to complete. The engineer has to create a CAD model by measuring physical part with calipers and/or a CMM and then create features and surfaces in model print is available, this data must be converted to define theCAD.afeatures.

This process is very prone to error and several iterations must be executedbefore an acceptable completed part is produced.There is now an available process to support "reverse engineering" applications which greatly reduces the time to complete these projects, provides a complete andaccurate CAD model on the first iteration, allows for the programlning of tool paths, and is completely compatible with the company's current CAD/CAM procedures. This process includes the integration of the automated 4-axis 3D laser digitizing from Digibotics, Inc. with a complete software environment for point processing, surface generation, and analysis from Imageware. A case study using these technologies was InterMotive recently completed by Fred Nicholas, CAD/CAM Manager, Technologies, a design and engineering company located in Belleville, MI. This paper describes the entire flow through the digitizing, surfacing, and CAD/CAM processes.ofFor this case study, Mr.

This paper describes the entire flow through the digitizing, surfacing, and CAD/CAM processes.ofFor this case study, Mr. Nicholas selected an automotive· thermostat housing which is a part with some complex outer surfaces and an internal cavity. A core stick of the internal cavity was created. The Digibot II laser digitizer from Digibotics, both the external part and the internal core stick. After editing Inc. was used to digitized data, a point file was imported into the Surfacer software system from the Imageware. Surfacer was used to create a network of B-spline curves and surfaces. NURBS surface file was created and then imported into several CAD/CAM Ansystems to demonstrate the ability to create a surface and tools paths for making the the Computervision CADDS5 software was used to part.

demonstrate the complete CAD/CAM process.For this case study,The Digibotics 3D laser digitizing technology was used for this project because it offers several unique advantages for "reverse engineering" of cOlnplex parts. The Digibotis an automated 4-axis digitizer that provides a simple, accurate, and quick way to copy or inspect complex sculpted surfaces. After fixturing an object, the user specifies a point density and instructs the system to begin. While digitizing, the object307rotates on the platter and the laser translates horizontally and vertically. Vertically spaced cross-sections are digitized starting from the bottom and moving upward. TheDigibot Adaptive Scanning Software intelligently follows the contours of the surface acquiring a sequence of adaptively spaced points while keeping the laser beam normal to the surface. Concave surfaces and multiple contours are obtained by pivoting the laser about obstructing surfaces.

Concave surfaces and multiple contours are obtained by pivoting the laser about obstructing surfaces. After digitizing, the object is composed of a stack of cross-sectional contours. The entire object along with individual points and contours can be examined and modified as needed using the Digibot Data Editor. A surface mesh composed of triangular facets is generated by connecting points and contours between cross-sections. The resulting polyline and/or surface mesh can be exported via data formats such as DXF, STL, VDA, or IGES.The Surfacer software system provides the link between the 3D laser digitized data and the conventional CAD/CAM system. Surfacer accepts point data directly from the Digibot Editor and performs point data analysis and processing such as segmentation, sectioning, filtering, and feature extraction. The NURBS surfaces are then quickly and accurately created from the selected point sections.

The NURBS surfaces are then quickly and accurately created from the selected point sections. The quality of the created surfaces are verified by points to surface cOlnparison, reflectance analysis, curvature analysis, and cross-sectional tools. These surfaces can then be edited and/or modified. Point, curve and/or surface geometry can then be exported directly to "downstream" processes such as CAD, CAM, Analysis or Visualization/Animation systems.The first step of the process was to digitize the thermostat housing using theDigibot II 3D laser digitizer. The housing was attached to the platter of the Digibot II by simply using hot glue. The Digibot II then digitized the entire housing by using Digibotics I patented adaptive scanning procedure. Both the external housing and the internal core stick were digitized. In a post-processing phase, Mr. Nicholas viewed the 120 cross-sections and performed minor editing to guarantee a good clean file of 3D data points.

In a post-processing phase, Mr. Nicholas viewed the 120 cross-sections and performed minor editing to guarantee a good clean file of 3D data points. The digitizing and editing process for this housing was completed in 1 1/2 days. Mr. Nicholas also used the Digibotics Triangulator and STL Generation softwareto create an STL file of the housing. The Accelerated Technologies Inc. DTM rapidprototyping system was used to create a polycarbonate copy of the housing. This isvery useful as an intermediate step to demonstrate that the digitizing process was verycomplete and accurate.The two files of 3D data points (containing about 20,000 points) were thenimported directly into Surfacer for creation of the surfaces. The two data files werefirst combined to form one complete external and internal object by registering the two308parts to a common axis through the centroid of the part.

The two data files werefirst combined to form one complete external and internal object by registering the two308parts to a common axis through the centroid of the part. Surfacer provides a completeset of tools for aligning multiple scans or aligning scan data to CAD data. A "push button" surface of the entire housing was then created. This one large complex NURBS surface is very useful for visualization, for gross product testing, and for interference checking with other CAD components. Because the Digibot digitized datais created in orderly well-spaced cross sections with no redundant or over-lapping data,this makes it much easier for Surfacer to create surfaces and perform its analysis functions. After the point analysis and registration, individual accurate surfaces were created. This occurs by generating surface patches over different point sections to define unique surfaces.

After the point analysis and registration, individual accurate surfaces were created. This occurs by generating surface patches over different point sections to define unique surfaces. Surfacer provides tools for detecting surface edges and segmenting points to define areas to be surfaced. Surface patches can then be fit to the points, with the user having control over the NURBS smoothness and fitting tolerances. For the housing, 21 different surface patches were created. These patches are stitched to create continuous, "watertight" connections, and different analysis was performed to confirm the accuracy of the surfaces. The 21 surface patches were then converted to an IGES 5.0 NURBS surface file for export into the CAD/CAM systems. This processing time in Surfacer took about two working days.Mr. Nicholas then imported the IGES file into CADDS5 for processing.

This processing time in Surfacer took about two working days.Mr. Nicholas then imported the IGES file into CADDS5 for processing. The file was also imported into ProEngineer and AutoCAD to demonstrate the IGES In CADDS5, a complete solid model was compatibility with other CAD systems. created by sewing the 21 individual surfaces. With this solid model, physical property analysis can then be performed such as volume and mass calculations. An STL file wascreated from this solid model, and a prototype was created using the 3D Systems rapid prototyping system. A tool path was created using Con1putervision I s CVNC software. The internal cavity was first isolated from the external part for tool path generation.Then a tool path was generated on the external surface. Finally, a product drawing was generated for documentation of this part.

Finally, a product drawing was generated for documentation of this part. The detailing environment included all radii, dimensions, cutting sections, tolerances, and information for placing the housing in its engine position. The procedures implemented in CADDS5 are very straight forward and utilizes the normal CAD/CAM functions that would be used for any solid model design. The procedures for CADDS5 were executed in two working days.The processes that are demonstrated in this case study were not available one year ago. There has been tremendous progress during this time in the integration of thedigitizing hardware/software and surfacing software for supporting these complexprojects. This case study was implemented in about one week, and is at least a four week improvement over uses of other approaches. The Digibotics and Imageware309week improvement over uses of other approaches.

The Digibotics and Imageware309week improvement over uses of other approaches. The Digibotics and Imageware systems are the only hardware/software combinations that provide automated digitizing and surfacing of complete complex objects. This capability to recreate 3D objects with this process is available now for applications such as creating spare parts, parts where no prints or CAD model exist, and creating a CAD model from a physical model done in clay, foam, or other fragile material..

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperApplications of 3D topography scanning and multi-material additive manufacturing for facial prosthesis development and productionMazher. I.

I. Mohammeda, Joseph Tatinenia, Brenton Caddb, Greg Peartb and Ian GibsonaaSchool of Engineering, Deakin University, Waurn Ponds Campus, Geelong, Victoria 3220, Australia bFacial Prosthetics, The Royal Melbourne Hospital, City Campus, 7 East, Main Block, Grattan Street, Parkville Victoria 3050Corresponding Author: mazher.mohammed@deakin.edu.auKeywords: Prosthesis, 3D printing, Modelling, Multi-material, Soft TissueAbstractProsthetic based rehabilitation offers several advantages over surgical intervention, however, devices are generally handmade using labour intensive and subjective manufacturing techniques. We investigate the use of optical scanning to capture the surface topography from a volunteer’s facial anatomy, reconstruct this into a 3D CAD model, and from that design a patient specific prosthesis. This approach offers many advantages over existing techniques as data collection is non-intrusive, rapid and provides anatomically precise information.

This approach offers many advantages over existing techniques as data collection is non-intrusive, rapid and provides anatomically precise information. A CAD approach affords greater flexibility when evaluating design iterations and allows for the creation of ‘parts libraries’ for use with patients with no initial reference anatomy. The final prosthesis is realised through high resolution, multi-material 3D printing for precise model reproduction and to add functionalities such as mimicry of soft and hard tissues. Ultimately, we believe our approach provides an optimised, low-cost approach for streamlining the complete methodology for prosthesis production.1 IntroductionIn a human population facial defects can arise as a result of congenital deformities, disease infiltration, and trauma. Given the prominence of the face and how it influences human interactions, such disfigurements can have a profoundly negative impact on the quality of life of a patient.

Given the prominence of the face and how it influences human interactions, such disfigurements can have a profoundly negative impact on the quality of life of a patient. With respect to rehabilitation, there are primarily two treatment option, which comprise surgical intervention or the use of a prosthesis. The decision making process as to which option is the most suitable is not so clearly defined and is dependent on a number of factors, including the size/severity of the condition, age, etiology, and critically, the patient's own personal preference [1, 2]. Examining prosthesis based rehabilitation, there are several immediate advantages when compared to surgical intervention such as the immediate aesthetic improvement, its simplicity over surgery, the ability to explore several design iterations without impact on the patient and its low cost.

More recently there has been a surge of interest in a1695tissue engineering approach to replace missing or compromised organs [3, 4], however despite the promise of this technology there are still many issues to resolve before this is likely to become a reality. Therefore prosthetic treatment provides a more robust, tried and tested approach which has a quick turnaround time for part production and does not have complications associated with surgical intervention, such as tissue rejection.Traditional prosthesis production is a long, labour intensive process, requiring the use of several invasive and subjective techniques over the entire fabrication process. A summary of the fabrication stages can be found in Figure 1. Typically the process begins with some form of casting approach, generally using plaster, to ascertain the topography of the defective area, or uncompromised anatomy that could be used as a template for the prosthesis [5, 6].

In some instances the plaster is placed over the entirety of a patients face, requiring breathing to be performed through a straw until the plaster sets. Consequently, due to the discomfort of this process the patients can often move during the casting process, resulting in an inaccurate topography map. Following the formation of plaster cast, a wax model is formed and manipulated to realise the finished prosthesis model. At this stage fixation device alignment is also performed, embedding either mechanical or magnetic abutments into the wax model. The process of alignment and finishing of the wax model are all performed manually, with the end result being wholly dependent on the artistic skill of the clinician, which is arguably subjective in nature [7]. Once the wax model is finalised, it is formed into a plaster negative, which in turn is used to form the final silicone prosthesis.

Once the wax model is finalised, it is formed into a plaster negative, which in turn is used to form the final silicone prosthesis. Any additional touches to the model, such as the addition of colours and hair, are again performed manually. This makes the whole process of prosthesis production a very arduous and labour intensive process.Figure 1: Traditional process chain for prosthesis fabrication1696Traditional prosthesis development is on a turning point, where several disruptive 3D technologies are likely to transform traditional prosthesis production. Modern optical scanning techniques readily allow for the rapid, high resolution reproduction of surface topographies, to a precision of <100µm while capturing useful data such as texture maps of a patients skin [8- 10]. Additionally, data is obtained non-invasively and can performed using laser free methodologies, greatly improving the potential uptake of this technology in a clinical capacity.

Additionally, data is obtained non-invasively and can performed using laser free methodologies, greatly improving the potential uptake of this technology in a clinical capacity. Modern Computer Aided Design (CAD) software readily allows for the conditioning of 3D scan data to create a complete model, with the added advantage of allowing for the digital storage of design iterations and the final model. Such digital data conditioning and storage, when applied to prosthesis production, compares very favourably to traditional hand crafting techniques. Finally modern 3D printing technologies offer the ability to easily and rapidly reproduce a high resolution digital model into a physical part [10-16].

Finally modern 3D printing technologies offer the ability to easily and rapidly reproduce a high resolution digital model into a physical part [10-16]. Fabrication can also be achieved using a vast array of flexible and biocompatible materials, allowing for augmentation of existing practises, and potentially, direct prosthesis fabrication.In this study we have investigated the use of optical scanning, CAD and 3D printing in the direct production of various facial prosthesis (ear and nose). Previous studies have primarily focused on the use of such technologies for the production of a cast to augment traditional techniques, or models which are made in rigid plastics. By contrast, the novelty in this work is to use high resolution 3D printing of flexible materials for the direct production of a prosthesis which not only is of a higher quality surface finish as compared to previous studies but also mimics the tactile feel and pigmentation of human soft tissues.

We also investigate the production of advanced prosthesis models, comprising multi-model and multi materials designs, which more closely reproduces human tissue through mimicry of both skin and cartilage. Our technique offers several advantages over traditional techniques as the use of optical scanning for topography mapping is non-invasive, can acquire data within minutes and realises anatomically precise, high resolution data, ideal for prosthesis production. A digital CAD approach for prosthesis design is superior when compared to traditional casting and hand crafting approaches as design iterations can be easily digitally stored, do not require any fabrication/material consumption and allows for operations such as mirroring of a model with relative ease.

Finally, the use of high precision 3D printing allows for rapid digital part realisation, can produce sophisticated prosthesis comprising complex multi-material/models and can readily allow for precise reproduction of a given model should duplicates be required. Ultimately, we believe the techniques presented in this work realise a patient specific, low-cost and high resolution approach to streamlining prosthesis optimisation and production.2 ExperimentalIn this study we have investigated the use of several 3D digitising, rendering and printing technologies to directly create prosthesis replica from a person’s anatomy.

The complete process chain can be seen in Figure 2, where surface topography maps are made using an optical scanner, designs are post processed using CAD and the final model is realised using high resolution, multi-material 3D printing.1697Figure 2: Diagram illustrating the proposed process chain for prosthesis production2.1 Topography mapping and model constructionIn this study a laser free, optical scanning system (Spider, Artec, Luxembourg) was employed to obtain the surface topography of a nose and ear of a volunteer. Laser-free scanning technology alleviated any health concerns resulting from laser exposure to the eyes. The scanner used had image scan resolution of approximately 50-100µm, which is more than adequate to resolve all the major and minor details of the anatomical part rendered. The scanner operates alongside a proprietary software (Artec studio 10, Artec, Luxembourg), which allow for the real-time visualisation of the scan data during acquisition.

The scanner operates alongside a proprietary software (Artec studio 10, Artec, Luxembourg), which allow for the real-time visualisation of the scan data during acquisition. It was found that several translations of the scanner were required to obtain the full surface map, comprising translations at approximately 10cm/s in a lateral and vertical arcing motion. The scanner allowed for a part to be rendered rapidly within approximately 3-5 minutes.The scanner software processes the input data as a point cloud, which is converted into a full contour map. Rudimentary operations can also be performed to condition the data, removing spurious noise, cropping of unrequired data, filling of holes/gaps in the model and smoothing of contours.

Rudimentary operations can also be performed to condition the data, removing spurious noise, cropping of unrequired data, filling of holes/gaps in the model and smoothing of contours. The resulting output is as an enclosed model, and so further post processing is required to hollow a part and realise features such as nasal cavities and ear canals.2.2 Model construction and Post processingData from the scanner was conditioned using more sophisticated CAD software to process the surface topography data and to construct more advanced design for composite model printing. In this study all additional post processing was performed using the 3-Matic software package (Materialise, Belgium), which can allow for direct STL manipulation, error1698checking and part thickness analysis.

Data conditioning comprised smoothing of rough surfaces, removal of features that could lead to print failure, and to perform procedures such as part hollowing.In this study we aimed to realise an advanced prosthesis that more closely mimics human physiology in terms of tactile feel and pigmentation. As a demonstrator we took a constructed model of the ear and reconstructed it to build separate models of the cartilage and a composite of the softer tissues. Facets of the cartilage model were constructed by a purely design based approach and so to obtain anatomical accuracy the model was constructed by cross referencing anatomical drawings of the ear cartilage and by direct feel of the test subjects actual ear. Approximations were then used to determine the layer thickness of the skin relative to the cartilage in the model.2.3 3D PrintingOnce rendering and post processing had been completed, the final models were directly 3D printed to produce the final prosthesis.

To obtain a high accuracy in digital reproduction a high resolution 3D printer was used (Connex 3 500, Stratasys, USA) which can print models in up to three individual materials or blends thereof to an accuracy of 16-30µm. This level of accuracy was more than adequate to reproduce all major and minor surface contours, whilst also realising a high quality surface finish to the prosthesis, similar to that obtained by traditional techniques. The printer operates exclusively using a set of materials which are developed by Statysys for use in this printer, with no option to use non-proprietary materials. These system is generally designed for product prototyping and design evaluation purposes and so facets relating to exact composition and biocompatibility of the materials used are not known.

These system is generally designed for product prototyping and design evaluation purposes and so facets relating to exact composition and biocompatibility of the materials used are not known. Exploration of the material properties will be the subject of our future work.Initially, a model is loaded into the printer’s software as an STL file and is allocated a specific material combination before being sliced into the individual printing layers. Simultaneously a water soluble support material is allocated as required to ensure the build integrity. The printer operates using a Polyjet technology, whereby a proprietary liquid photo curable polymer is delivered by the print head and subsequently flattened and cured by a UV lamp. Once a model is complete, a final cleaning phase is required to remove the support material before the part is ready for use.

Once a model is complete, a final cleaning phase is required to remove the support material before the part is ready for use. The printing process can take on average 2-3 hours for a part to be printed, depending on its size and orientation.The printer used in this study was capable of printing in several different materials, which have a variety of different colour and mechanical properties. For instance the Tango plus material range is a flexible material, while the Vero materials are rigid coloured materials. For the final prosthesis parts we examined a combination of both Tango plus and Vero materials such that we could obtain a final model with a flexible tactile feel but also with adjustable colours. It was our hope in this study to realise combinations that mimic the feel and pigmentation of a person’s actual tissue.

It was our hope in this study to realise combinations that mimic the feel and pigmentation of a person’s actual tissue. Further, the advanced multi model approach allows for further control of the tactile feel of a given anatomical part.16993 Results3.1 Scanning and model creationVarious scans were performed on a test subject in an attempt to reproduce the surface topography of their nose and left ear. In this instance the subject was a health individual who suffered no facial defects, allowing for comparison of the printed prosthesis with the original anatomy. When performing scans, the fast fusion mode of the scanner was utilised, which allowed for rapid, real-time visualisation of the scans as they were being performed. Scans were performed translating the scanner through the various orientations previously described.

Scans were performed translating the scanner through the various orientations previously described. Following completion of the data acquisition several additional data conditioning phases were performed using Artec studio to remove spurious nose and to reconstruct gaps in the data. Following completion of the scans, the software allowed for several automated procedures to crop the part of interest from the wider facial data and to convert the scan from a surface to an enclosed mesh.It was found that there were limitations to the scanning process and areas of the skin which were shiny/reflective were difficult to render during scanning. This limitation could be overcome using a mat finish power to dampen optical reflections, however, that was not required in this study. Additionally, regions that had low levels of light exposure were equally difficult to render, such as the nasal cavities and ear canal.

Additionally, regions that had low levels of light exposure were equally difficult to render, such as the nasal cavities and ear canal. As such features are critical to the final prosthesis, they had to be rendered independently using 3-matic.3.2 Model post processing3.2.1 Single modelFollowing initial rendering, the models were checked to remove errors such as inverted normals, multiple shells, noisy shells, etc. This procedure ensure the best quality of digital data to be used for the subsequent design phases. In general the rendered models from the scanner looked reasonably close to the final prosthesis models and so only relatively simple post processing was required. With respect to the nose data, the nasal cavities were completely enclosed and so the first procedure was to remove excess digital material to open these areas and to provide access to the reverse side of the model.

The necessity for this is that generally a recipient of such a prosthesis will have use of their nasal ducts and so open access here would allow for a potential patient to retain the ability to smell, breathe, etc. With respect to the ear data, by cross referencing the major contours against the original test subject, it was found that several of the contours formed by the cartilage had been lost during the post processing within the Artec Software. Therefore these areas were manually reconstructed using the ‘push pull’ and extrude functions within 3-matic. Figure 3a) illustrates the final model of the ear prosthesisWith respect to the ear, when a thickness analysis was performed, it was found that there were regions which were <350µm thick. While this is not an issue to the aesthetics of the digital mode, should this be printed, this region would be extremely fragile.

While this is not an issue to the aesthetics of the digital mode, should this be printed, this region would be extremely fragile. From initial test prints at this thickness in the rubber like Tango plus material, it was found that the parts 1700ruptured during the support material removal phases. Several design iterations were examined increasing the minimum thickness in steps of approximately 200µm, where it was found that for thicknesses greater than 1mm the models could be cleaned without rupturing. Therefore the thickness analysis of the part is critical to the integrity of the final prosthesis using the Tango plus material, and the parts were thoroughly inspected to ensure no regions were of a thickness less than 1mm. Figure 3b) shows a model of the ear before and after the thickness analysis and subsequent model region growing.Figure 3: a) Various orientations of the ear prosthesis model.

Figure 3b) shows a model of the ear before and after the thickness analysis and subsequent model region growing.Figure 3: a) Various orientations of the ear prosthesis model. b) Colour map thickness analysis of the ear model for i) the raw data model input from the 3D scanner software, ii) post processed model to grow the thickness of the inner ear section to approximately ≥1mm.3.2.2 Multi-component ear modellingThe anatomy of the human ear comprises a single piece of cartilage tissue in the ear which makes the design process relatively simple in comparison to more complex cartilage structures such as the nose. Initially attempts were made to render the cartilage of the ear as a stand-alone model. To achieve this the original ear model was duplicated and reduced in size1701by 96% to create a replica structure that was positioned approximately 1.5mm into the original ear model.

To achieve this the original ear model was duplicated and reduced in size1701by 96% to create a replica structure that was positioned approximately 1.5mm into the original ear model. This new model was then reformed by referencing anatomical drawings of a generic human ear cartilage, the contours captured within the scanned ear model and by direct feel of the human subject. The cartilage reconstruction process can be performed to a high degree of precision using digital thickness analysis to augment the design process, however arguably the use of feel to determine model topography is a subjective metric. It is noted that this part of the process is merely for qualitative assessment of the model and forms a very minor part in the design process and does not compromise the precision of the final rendering.

The final model that was achieved can be seen in figure 4, which closely matched the medical diagrams.Figure 4: a) A classical medical diagram of the cartilage within the ear [17], adapted from https://lookfordiagnosis.com/mesh_info.php?term=ear+cartilage&lang=1. b) The multi-model design of the ear prosthesis comprising the soft tissue composite and the cartilage. For visualisation purposes both segments are also represented individually.3.3 Prosthesis 3D printing3.3.1 Single model printingFollowing completion of the models, initial tests were performed to ascertain the printing precision, model integrity, mechanical conformity and pigmentations that could be achieved using a multi-material combinatory approach. The Connex printer is capable of printing with 3 materials simultaneously. However as flexibility is achieved through the use of the transparent Tango plus material, only two additional materials (colours) could be used.

However as flexibility is achieved through the use of the transparent Tango plus material, only two additional materials (colours) could be used. Given the availability of colours by the manufacturers, it was decided that blends of magenta and yellow would provide the best options for skin pigmentation mimicry. An effective colour map was produced of the different material combinations when using Tango plus and Vero materials as can be seen in figure 5a). Another issue that arose was the percentage blend of the rigid Vero with the flexible Tango plus material, where a threshold of >50-60% Vero material 1702compromised the tactile feel beyond the desired flexibility found in a typical prosthesis. The tango plus material used in this study was a translucent variant, which on its own provided the softest tactile feel but was not a suitable colour for a prosthesis.

The tango plus material used in this study was a translucent variant, which on its own provided the softest tactile feel but was not a suitable colour for a prosthesis. It was found that a minimum of 10% Vero material was required to provide any visually noticeable pigmentation into a given prosthesis. Various combinations of Tango and Vero material were explored to print the final prosthesis, and the various material combinations examined are illustrated in Figure 5a).Figure 5: a) i) An image of the material map which show the various colour combinations possible with the 3D printer using two Vero colours and the flexible Tango plus material, and images of various printed prosthesis of ii) an ear and iii) a nose model. Material combinations are also highlighted in the legend.

Material combinations are also highlighted in the legend. b) Comparative images of the subject’s ear and the 3D printed model from i) side and ii) reverse profiles.The high resolution of the connex printer and it effective use of dissolvable support material allows for the reproduction of both the ear and nose models and with a smooth surface finish. The reproduced models were found to be a very good match to the original anatomy,1703both in terms of major surface contours and overall size. Various nose and ear models can be seen in figure 5aiii), where is can be clearly seen how the multi-material printing allowed for a diverse array of prosthesis pigmentations from lighter to darker skin tones. Also show in figure 5b) is a comparison of the printed ear prosthesis relative to the volunteer subject’s original ear. It can be seen that the final prosthesis very closely matches the original contours of the ear, validating the efficacy of this technique.

It can be seen that the final prosthesis very closely matches the original contours of the ear, validating the efficacy of this technique. All 3D printed models were realised using a single material combination across the entirety of the model. When examining the ear models more closely, it was found that the richness of the visible colour changed depending upon the thickness of the material. By comparison the nose models did not exhibit such features. This is perhaps due to the thickness tolerances of these models, which were generally between 1 and 6mm allowing for varying degrees of translucency. In contrast the minimum thickness of the nose was approximately 15mm therefore at these thicknesses all translucency properties are negligible. This facet is similar to human physiology, where the thickness of the soft tissues overlaying the cartilage of the ear has a variety of different pigmentations owing to its relative thickness.

This facet is similar to human physiology, where the thickness of the soft tissues overlaying the cartilage of the ear has a variety of different pigmentations owing to its relative thickness. We hope to examine such facets more closely in future work.The printed models were qualitatively assessed for their tactile feel and mechanical compliance to that or the original human subject. Tests comprised the ability to flex the lobe area, a vertical compression test from the lobe to the upper portion of the ear and a lateral compression test pinching across the centre of the ear. It was found that the printed models all exhibited similar characteristic to the original anatomy, such that upon relaxation of the compressive force that the ears would return to the neutral position of the model, as can be seen in figure 6a). It was noted that the printed material was more rigid, despite its elasticity, than real human anatomy.

It was noted that the printed material was more rigid, despite its elasticity, than real human anatomy. This resulted in a greater force having to be applied to achieve the various modes of compression, highlighting limitations of current printable, flexible materials. The purpose of this work was to present our methodology regarding the scanning, design and printing capabilities and we hope to more quantitatively measure mechanical properties of the prosthesis in future studies. It is also noted that as the percentage combination of the Vero material was increased in the model, the necessary applied force became larger. At a percentage of 60% Vero, the elasticity of the model was compromised beyond an acceptable level and became noticeably rigid. This therefore limits the wider colour combinations possible while still retaining acceptable levels of elasticity to mimic human mechanical properties.

This therefore limits the wider colour combinations possible while still retaining acceptable levels of elasticity to mimic human mechanical properties. We hope in future work to quantify such forces both for real anatomy and printed prosthesis.3.3.2 Multi-model printingBeyond a single material model, we attempted to realise multiple models, rendered using independent material combinations, thereby mimicking the softer and harder tissues of the ear. The connex printer allows for the overlay of multiple models and for independent material allocation to each when printing. We therefore processed the advanced ear model such that the soft tissue composite was printed with 100% Tango plus and the Cartilage model was printed with 50% Tango plus and 50% VeroMagenta. These material combinations were used primarily for visualisation purposes such that each element could be visually differentiated,1704and the final printed model can be seen in figure 6b).

These material combinations were used primarily for visualisation purposes such that each element could be visually differentiated,1704and the final printed model can be seen in figure 6b). The connex printer was found to provide excellent multi-material printability, with the two materials seamlessly blending into a single structure without any impact on the final surface finish of the model. Therefore the polyjet printing process of the connex is considered ideal for rendering of blended materials as well as multiple models for direct prosthesis printing.Once again the mechanical properties of the model were assessed by qualitative compression tests. On this occasion it was found that there was much greater rigidity to the ear model, as expected, however the ability to be compressed vertically and laterally were not compromised and the model would also return to the neutral position upon relaxation of the compressive force.

It was noted that the movement of the ear lobe was identical to the single material/model prints. Outcomes of the multi-model printing illustrate that the mechanical properties of the printed models can be further modified to reach increasing levels of complexity, as found in actual human anatomy. We hope in the future to more quantitatively assess the mechanical properties and to explore the ability to more seamlessly blend the colour combinations into a more realistic final modelFigure 6: a) Qualitative compression of the model to demonstrate the realism in the tactile feel for both i) vertical and ii) lateral compression. b) Multi-model printing of the ear, with independent material combinations for the cartilage and residual soft tissue composite.17054 ConclusionsThis study has demonstrated the potential of the 3D design and multi model/material printing, augmented with the use of optical surface scanning, to produce realistic prosthetic models of both the ear and nose.

The fabricated prosthesis were realised to a high degree of accuracy and surface finish, and we believe the technique suitable to render additional prosthetic parts, such as orbital prosthesis. We realised advanced prosthesis models beyond the traditional single model variants using novel design techniques to render components such as cartilage alongside other soft tissues (skin, etc). Using the multi-material printing approach we could tailor the skin pigmentation of the prosthesis to a variety of skin tones, whilst also mimicking the mechanical properties of the original anatomy. The mechanical properties can be further tailored using a multi-model approach. Currently there are limitations in the complexity of the skin tones that can be mimicked without compromising the mechanical properties, due to the percentage material combinations. Additionally, the overall printed prosthesis tactile feel is more rigid than actual anatomy.

Additionally, the overall printed prosthesis tactile feel is more rigid than actual anatomy. These limitations are primarily due to the materials used in the 3D printer, but we believe as the technology matures over the coming years, these limitations will be resolved. Ultimately, the findings in the work validate our approach for direct prosthesis production which overcome limitations relating to the subjective nature of current prosthesis fabrication, and allow for production within a single day. This compares favourably to traditional techniques where typically a prosthesis if fabricated over several weeks/months.

This compares favourably to traditional techniques where typically a prosthesis if fabricated over several weeks/months. Ultimately, technique holds considerable potential for implementation within a clinical setting, streamlining to overall process for prosthesis production and could see applications in other niche areas such as soft robotics or anatomical modelling.thisAcknowledgementsWe would like to thank the Royal Melbourne hospital for their input on the clinical aspects of this project. We would also like to thank the School of Engineering at Deakin University who provided funds and resources for this pilot project and their technical staff who assisted in the 3D printing of the models.References1.2.3.4.Kretlow, J.D., A.J. McKnight, and S.A. Izaddoost, Facial Soft Tissue Trauma. Seminars in Plastic Surgery, 2010. 24(4): p. 348-356. Ranganath, K. and H.R. Hemanth Kumar, The Correction of Post-Traumatic Pan Facial Residual Deformity.

plications of a Hybrid Manufacturing Process for Fabrication and Repair of Metallic StructuresFrank Liou*, Kevin Slattery**, Mary Kinsella+, Joseph Newkirk*, Hsin-Nan Chou**, and Robert Landers*University of Missouri – Rolla, Rolla, MO, 65409 ** The Boeing Company + AFRL/MLLMPAbstract Since its appearance, rapid prototyping technology has been of interest to various industries that are looking for a process to produce/build a part directly from a CAD model in a short time. Among them, the direct metal deposition process is the only process which directly manufactures a fully dense metal part without intermediate steps. However, challenges of the direct metal deposition process include building overhang structures, producing precision surfaces, and making parts with complex structures. Coupled between the additive and the subtractive processes into a single workstation, the integrated process, or hybrid process, can produce a metal part with machining accuracy and surface finish.

Coupled between the additive and the subtractive processes into a single workstation, the integrated process, or hybrid process, can produce a metal part with machining accuracy and surface finish. Therefore, the hybrid process is potentially a very competitive process to fabricate and repair metallic structures. This paper summarizes the current development of the hybrid process to process high temperature metallic materials, including tool steel and Ti64. Research in simulation and modeling, process development, and actual part building and repair are discussed.Introduction For more than a decade, layered manufacturing technology, also known as Rapid Prototyping (RP) has given industry an approach to achieve the goal of providing products in a shorter time and at a lower cost. Most of the current RP systems are built on a 2.5-D platform.

Most of the current RP systems are built on a 2.5-D platform. Among them, the laser-based deposition process is a potential technique that can produce fully functional parts directly from a CAD system and eliminate the need for intermediate steps. However, such a process is currently limited by the need for supporting structures – a technology commonly used in all the current RP systems. Support structures are not desirable for high strength and high temperature materials such as metals and ceramics since these support structures are very difficult to remove. Multi-axis systems can offer much more flexibility in building complex objects. Laser aided RP is advancing the state-of-the-art in fabrication of complex, near–net shape functional metal parts by extending the laser cladding concept to RP.

Laser aided RP is advancing the state-of-the-art in fabrication of complex, near–net shape functional metal parts by extending the laser cladding concept to RP. The Laser Aided Manufacturing Process being developed in the Laser Aided Manufacturing Processes (LAMP) Laboratory at the University of Missouri-Rolla (UMR) combines laser deposition and machining processes to develop a hybrid rapid manufacturing process to build functional metal parts. This paper summarizes the research and applications of such a hybrid process for fabrication and repair of metallic structures.A Hybrid Manufacturing System In order to expand the applications of metal deposition processes, multi-axis capability is greatly needed. A multi-axis rapid manufacturing system can be hardware-wise configured by adding extra degrees of mobility to a deposition system or by mounting a laser deposition device on a1multi-axis robot.

A multi-axis rapid manufacturing system can be hardware-wise configured by adding extra degrees of mobility to a deposition system or by mounting a laser deposition device on a1multi-axis robot. The configuration could also be a hybrid system in which a metal deposition system is mounted on a multi-axis CNC machine. With the addition of extra rotations, the support structures may not be necessary for the deposition process in order to build a complicated shape. Figure 1 illustrates the process to build an overhang structure on a 2.5D and multi-axis deposition system. Due to the nature of the deposition process, it is driven by a so- called “slicing” procedure, which uses a set of parallel planes to cut the object to obtain a series of slicing layers.

Due to the nature of the deposition process, it is driven by a so- called “slicing” procedure, which uses a set of parallel planes to cut the object to obtain a series of slicing layers. So far, the slicing software on the market is only able to handle 2.5D slicing in which the building/slicing direction is kept unchanged (usually Z+ direction) and it lacks the capability of changing directions to fully explore the capability of multiple degrees of freedom.Part to be builtSupport structureDeposition systemRotation(b) Fig 1. (a) build part with support structure; (b) with multi-axis capability, after building the column, the table can be rotated; (c) After rotation, continue to build the component from another direction.(a)(c)This process uses laser deposition for material deposition and CNC milling for material removal. It includes two major systems: a laser deposition system (Rofin–Sinar 025) and a CNC milling machine system (Fadal VMC-3016L).

It includes two major systems: a laser deposition system (Rofin–Sinar 025) and a CNC milling machine system (Fadal VMC-3016L). The laser deposition system and CNC milling machine work in shifts in a five-axis motion mode. The laser deposition system consists of a laser and a powder feeder. In conventional 2.5-D laser deposition process to create three-dimensional parts, overhang and top surfaces of hollow parts need to be supported. Often support materials for functional metal parts are not feasible. Moreover, deposition of the support material for metals leads to poor surface quality at the regions in contact with the support structure. Moreover, it increases the build time of the part and necessitates a time-consuming post-processing. Additionally use of support increases the build time of the part and necessitates a time- consuming post-processing. With a five-axis deposition process integrated with five-axis machining, these obstacles can be removed.

With a five-axis deposition process integrated with five-axis machining, these obstacles can be removed. This paper summarizes the issues and related approaches in the research and applications of the hybrid deposition-machining process, including laser deposition, sensing and control, and process planning.Laser Deposition Research Laser deposition is a solid freeform fabrication process that has the capability of direct fabrication of metal parts. The process uses a focused laser beam as a heat source to create a molten pool on an underlying substrate. Powder material is then injected into the molten pool through nozzles. The incoming powder is metallurgically bonded with the substrate upon solidification. The part is fabricated in a layer by layer manner in a shape that is dictated by the CAD solid model, which is sliced into thin layers orthogonal to the z-axis.

The part is fabricated in a layer by layer manner in a shape that is dictated by the CAD solid model, which is sliced into thin layers orthogonal to the z-axis. After the slice data is2then translated into laser scanning paths, an outline of each feature of the layer is generated and then the cross section is filled using a rastering technique to fabricate a single layer. After the deposition of a single layer, the subsequent layers are deposited by incrementing the nozzles and the focusing lens of the laser in the z-direction, until the three-dimensional part is completed. To successfully deposit powder material, in the laser deposition process, the powder beam and laser beam need to be carefully understood and controlled. Simulations of the phenomenon which occur during the laser deposition process are completed to increase knowledge of the process and measurements through sensors are conducted to validate those simulations.

Research and modeling of the powder delivery system and laser-powder interaction process have been conducted.The quality and efficiency of laser aided direct metal deposition largely depends on the powder stream structure below the nozzle. Numerical modeling of the powder concentration distribution is complex due to the complex phenomena involved in the two-phase turbulence flow. In this research, the gravity-driven powder flow is studied along with powder properties, nozzle geometries, and shielding gas settings [Pan05]. A 2-D non-spherical model includes (cid:533) and R, as shown in Figure 2 and was used for powder simulation. In that model, the parameter (cid:533) indicates how much the contact point C deviates from the foot of a vertical reference line through the particle gravity center and R shows the actual distance between the contact point and the gravity center. A random value of (cid:533) can be used to model irregular bouncing.

A random value of (cid:533) can be used to model irregular bouncing. Realistic simulation normally requires the extension of 2-D models to 3-D.rFig 2. A 2-D non-spherical model was used to model powder flow.(cid:550)VR(cid:533)C(a) (b)A 3-D numerical model was used to quantitatively predict the powder stream concentration variation in order to facilitate coaxial nozzle design optimizations as shown in Figure 3. A commercial computational fluid dynamic (CFD) code, FLUENT, was used to solve the momentum and turbulence equations to obtain the flow field and predict particle phase motions. Effects of outer shielding gas directions, inner/outer shielding gas flow rate, powder passage directions and opening width on the structure of the powder stream have been systematically studied.

Effects of outer shielding gas directions, inner/outer shielding gas flow rate, powder passage directions and opening width on the structure of the powder stream have been systematically studied. A series of experimental studies is designed to determine the effects of the outer gas injection angle and outer gas flow rate on the particle concentration mode as well as to validate the simulation. Measurements are carried out at 30°Fig 3. A 3-D simulation model to quantitatively predict the powder stream concentration variation for nozzle design and optimization.3and 60° outer gas settings respectively. In each setting, the outer gas flow rate is varied for 8.5, 13.7 and 22.1 litres/min, while keeping the inner gas flow rate constant at 2.43 litres/min.

In each setting, the outer gas flow rate is varied for 8.5, 13.7 and 22.1 litres/min, while keeping the inner gas flow rate constant at 2.43 litres/min. These experimental conditions can be summarized in Table 1:Case 1Case 2Case 3Case 4Case 5Case 6Case 7Nozzle angle30°/60°30°30°30°60°60°60°Inner gas (L/min)02.432.432.432.432.432.43Outer gas (L/min)08.513.722.18.513.722.1Table 1. Gas flow rates and directions used in experiments.The numerical simulation results are compared to the experimental data using prototyped coaxial nozzles as shown in Figure 4. The results are found to match and then validate the simulation. This study shows that the particle concentration mode is influenced significantly by nozzle geometries and gas settings.Fig. 4. Measured and calculated particle concentration variation along central axis.

4. Measured and calculated particle concentration variation along central axis. (a) case 2; (b) case 5; (c) case 3; (d) case 6; (e) case 4; and (f) case 7.4A melt pool formed during laser deposition is a critical factor and melt pool geometry is a crucial factor in determining clad quality. To obtain a high quality resulting part, a deep understanding of the underlying mechanisms is required. In this research, a mathematical model, as shown in Figures 5 and 6, was developed to simulate the coaxial laser cladding process with powder injection, which includes laser-substrate, laser-powder and powder-substrate interactions [Han04]. The model considers most of the associated phenomena, such as melting, solidification, evaporation, evolution of the free surface and powder injection.

The model considers most of the associated phenomena, such as melting, solidification, evaporation, evolution of the free surface and powder injection. The fluid flow in the melt pool, which is mainly driven by the Marangoni shear stress as well as particle impinging, together with the energy balances at the liquid-vapor and the solid-liquid interfaces are investigated. Powder heating and laser power attenuation due to the powder cloud are incorporated into the model in the calculation of the temperature distribution. The influences of the powder injection on the melt pool shape, penetration, and flow pattern are predicted through the comparison for the cases with powder injection and without powder injection. Dynamic behavior of the melt pool and the formation of the clad are simulated.

Dynamic behavior of the melt pool and the formation of the clad are simulated. The effects of the process parameters on the melt pool dimension and peak temperature are further investigated based on the validated model.Shielding GasLaser BeamPowderDeposited LayerRemelted Zone SubstrateMelt poolFig. 5. Schematic diagram of the calculation domain for laser cladding process.Interface0(cid:32)(cid:73)Fluid 1 nˆ0(cid:31)(cid:73)tˆTransition regionFluid 20(cid:33)(cid:73)Fig. 6. Interface transition region.5In order to give the quantitative comparison, the clad height is calibrated accurately for a length of 5.5 mm starting from the deposition point, and the resulting comparison is presented in Figure 7. As seen in the plot, the simulation result agrees well with the experiment and the error of the average height is around 8.7 %.

As seen in the plot, the simulation result agrees well with the experiment and the error of the average height is around 8.7 %. Four typical laser power levels, which are 300 W, 500 W, 750 W and 1000 W, were tested and the results are presented in Figures 8 and 9. It can be seen from the plot that although there is about 90 K over-prediction on average in peak temperature, the general trend between the simulation and experiment is consistent.0.45)m m0.40.350.3Fig. 7. Comparison of the clad height between simulation and experiment.(t h g i e H d a l C0.250.20.150.10.05Simulation Experiment00123456X (mm)3000)K(e r u t a r e p m e t25002000Exp erimentSimulationFig. 8.

8. Melt pool peak temperature comparison between simulation and experiment at different laser power level.k a e pl o o p1500t l eM10002004006008001000Laser p ower level (W)1.5Experiment1.2Simulation)m m(h t g n e ll o o p0.90.6Fig. 9. Comparison of melt pool geometry (length) between simulation and experiment for different power levels.t l eM0.30 2004006008001000Laser p ower level (W)Automated Process Planning System Process planning, simulation, and tool path generation for the LAMP allow the designer to visualize and perform the part fabrication from a desktop. The Laser Aided Manufacturing Process Planning uses STL models as input and generates a description that specifies contents6and sequences of operations. The objective of the process planning is to integrate the five-axis motion and deposition-machining hybrid processes.

The objective of the process planning is to integrate the five-axis motion and deposition-machining hybrid processes. The results consist of the subpart information and the build/machining sequence [Ruan05]. Basic planning steps involve determining the base face, extracting the skeleton, decomposing a part into subparts, determining build sequence and direction for subparts, checking the feasibility of the build sequence and direction for the machining process, and optimization of the deposition and machining.1) Skeleton Computation An algorithm for computing the skeleton of a 3-D polyhedron is needed. The algorithm is based on a classification scheme for points on the skeleton computation in which the continuous representation of the medial axis is generated with associated radius functions.

The algorithm is based on a classification scheme for points on the skeleton computation in which the continuous representation of the medial axis is generated with associated radius functions. Because it is used as a geometric abstraction, the skeleton is trimmed from the facets that touch the boundary of the object along every boundary edge for which the interior wedge angle is less than (cid:83) rad.2) Part Orientation The determination of the base face from which the building process of the part starts is very important. The base face functions as the fixture in the machining process. Therefore, when in the machining process, it must provide enough resistance against the cutting force. The maximal resistance force depends on the area of the base face.

Therefore, when in the machining process, it must provide enough resistance against the cutting force. The maximal resistance force depends on the area of the base face. The base faces have to satisfy the following conditions: 1) Located on the convex hull of the part, and 2) Certain amount of contact area.3) Part Decomposition and Building Direction The objective of part decomposition is to divide the part into a set of subparts, which can be deposited and machined. The topology of the part can be obtained from the skeleton. Each branch of the skeleton corresponds to a subpart. One of the partitions that is preformed is along a non-planar surface. Therefore, close to the partition area, 3-D layers are needed to build the connection between two subparts. The build direction of a subpart may not be constant. It changes when the part is built layer by layer so that for two adjacent layers, the later layer can be deposited based on the early layer without any support structures.

It changes when the part is built layer by layer so that for two adjacent layers, the later layer can be deposited based on the early layer without any support structures. To achieve the non-support build, the build directions need to be along the skeleton.4) Building Sequence The results of decomposition are recorded in an adjacency graph where nodes represent subparts, and edges represent the adjacency relationship between connected nodes. After considering part building order, a directed graph that represents the precedence relationship among subparts can be constructed. From the precedence graph, one can identify in what order the subparts can be built. With the precedence graph, a set of alternative building plans can be generated.

From the precedence graph, one can identify in what order the subparts can be built. With the precedence graph, a set of alternative building plans can be generated. Each plan represents a possible building sequence on the decomposed geometry and can be chosen optimally depending upon machine availability or other criteria such as minimum building time.5) Machinability Check The main purpose of a machinability check is to choose an optimal building sequence from the sequence set. Local and global collision checks are operated first to choose acceptable sequences since the building direction is different in each sequence. If any kind of collision happens or an under cut plane appears, the corresponding sequence will be discarded. For the rest of the7building sequences in the set, the buildability check and machining time computation is performed to find an optimal building sequence.Figure 10 shows a bearing seat example. The part was first computed to find the skeleton diagram.

The part was first computed to find the skeleton diagram. Based on the skeleton diagram, it is then decomposed into sub-components for slicing.(a)(b)(c)(d)Fig. 10. Example of a bearing seat.Implementation and Applications The part building process is demonstrated by several examples. Figure 11 illustrates a 3-D layer building process. In Figure 11 (b), a deposition result is shown. The STL model and process planning results are shown in Figure 11 (a). The side milling and top milling processes are demonstrated in Figure 11 (c) and (d) respectively. The final part with a total degree of 45 after conducting the final machining process is shown in Figure 11 (e). Compared to regular 2.5D rapid systems, the LAMP system saves about 80% of the support structures as shown in Figure 11.8Machining toolpathDeposition toolpathNote: the figure is scaled to show toolpath.(a) Process planning result(c) Side milling(d) Top milling(e) Final overhang partFig. 11.

11. A part built with 3-D layers.One unique application of such a hybrid deposition and machining system is for repair applications. In the LAMP process, the paths for deposition were generated in the software written in Visual C++ using the ACIS 3D toolkit. Layer height was set at 0.03 inches. The track width was set to be 0.1 inch with 50% overlap. The contour-parallel was used. The results are shown in Figure 12. The deposition paths were then sent to a postprocessor to generate the NC codes. Experiments were done for the third repair option as well. The result of the repaired part is shown in Figure 13 [Eiamsa-ard 05].9(a)(b)(c)Fig. 12: (a) Slices (b) Deposition paths (c) Deposition paths (close up)0.5F0.5F40mmd Th20mmbl=10mm10mmFig. 13: Part after repairing.Fig. 14: Bending test set up.The interfacial strength is determined from the four-point bend test (Figure 14).

13: Part after repairing.Fig. 14: Bending test set up.The interfacial strength is determined from the four-point bend test (Figure 14). The four-point flexture test is based on the storage of elastic energy on bending. Interfacial cracks propagate when the strain energy release rate equals to the critical energy release rate (Gc) of the interfacial failure. The four-point bend test has been used to analyze the interface between the substrate and the cladding produced by laser processing. The 50x6x1 mm specimens are cut out from the deposition. A center pre-crack is made on the specimen in order to induce symmetrical cracks along the clad-substrate interface. The specimen is then loaded in a four-point flexture on an Instron TT-B Universal Testing machine until a new crack propagates through the entire cladding. The interfacial fracture energy of the laser cladding tool steel specimen is compared to the tool steel weld specimen of the exact same dimension.

The modeling and simulation of powder flow and powder-laser interaction help design and set the process parameters for laser deposition. The overall goal for process planning is not only to find a solution to build a part but also to look for an answer to produce it in the least amount of time; therefore, the least amount of switching between the machining process and deposition process, the better since each switch requires retreating and relocating the deposition nozzle as well as the machining tool, which may cost extra time. The process planning analyzes the slicing results and compares them to the original geometry in STL format to find an optimal process sequence. With integration of multi-axis deposition and machining processes on the same work station, a hybrid system is able to produce complicated geometry, especially the overhang structure with less or no support structures, which saves material cost.

The switch between the machining and deposition process is optimized to save operation time/cost. Based on different geometry shapes, usually the LAMP system can save up to 50~60% of support structures. The surface quality of the final product is similar to the industrial milling capability.Acknowledgments This research was supported by the National Science Foundation Grant Number DMI-9871185, the grant from the U.S. Air Force Research Laboratory contract # FA8650-04-C-5704, and UMR Intelligent Systems Center. Their support is greatly appreciated.References • Eiamsa-ard, Kannayut, Hari Janardanan Nair, Lan Ren, Jianzhong Ruan, Todd Sparks, Frank Liou, ”Part Repair Using a Hybrid Manufacturing System,” Proceedings of the Sixteenth Annual Solid Freeform Fabrication Symposium, Austin, Texas, August 1-3, 2005.Han, L., F.W. Liou, and K.M. Phatak, “Modeling of Laser Cladding with Powder Injection,” Metallurgical and Materials Transactions B, 2005, VOLUME 35B, DECEMBER 2004, pp.

Liou, and K.M. Phatak, “Modeling of Laser Cladding with Powder Injection,” Metallurgical and Materials Transactions B, 2005, VOLUME 35B, DECEMBER 2004, pp. 1139-1150.Liou, Frank, Robert Landers, J. Choi, Sanjeev Agarwal, Vijayakumar Janardhan, S.N. Balakrishnan, “Research and Development of a Hybrid Rapid Manufacturing Process,” Proceedings of the Twelfth Annual Solid Freeform Fabrication Symposium, Austin, TX, August 6-8, 2001, pp. 138-145.Pan, Heng and Frank Liou, “Numerical Simulation of Metallic Powder Flow in a Coaxial Nozzle for the Laser Aided Deposition Process,” Journal of Materials Processing Tech., Volume 168, Issue 2 , 30 September 2005, Pages 230-244Ruan, J., K. Eiamsa-ard, and F. Liou, “Automatic Process Planning and Toolpath Generation of a Multi-Axis Hybrid Manufacturing System,” SME Journal of Manufacturing Processes, Vol 7, No. 1, pp. 57-68, 2005.11.

e Applications of Additive Manufacturing Technologies in Cyber-Enabled Manufacturing SystemsSriram Praneeth Isanaka and Frank LiouMissouri University of Science and TechnologyRolla, MO 65409AbstractThe application of networked sensors and control in various areas, such as smart grids and infrastructures, has become a recent trend, called cyber-physical systems. The Cyber Enabled Manufacturing (CEM) environment is to apply these technologies in manufacturing systems to handle a significantly greater magnitude of manufacturing data. Additive manufacturing techniques print or place material layer by layer to form a part, thus have a great potential to help accelerate CEM process by printing or embedding sensors and actuators in the proper locations. This paper summarizes the roles of additive manufacturing technologies to help establish a CEM environment.1.

This paper summarizes the roles of additive manufacturing technologies to help establish a CEM environment.1. Introduction1.1.Cyber-Enabled Manufacturing (CEM)A CEM system is an offshoot of the generic Cyber Physical System (CPS) which consists of embedded and distributed sensors that are networked over a wide area to effectively gather information while being coupled with control systems that offer immediate response. Their applications are acknowledged to include electricity control grids, traffic management and accident reduction, disaster prediction and response, and preventive maintenance systems [1]. Their application in a manufacturing environment leads us to the concept of CEM.A CEM system can be defined as an environment wherein intelligent use of networked sensors and process management is combined to efficiently utilize manufacturing information to increase awareness and connectivity [2].

An efficient CEM system would increase productivity, and safety standards while simultaneously driving down costs. A CEM environment enables remote access and control of manufacturing systems with the aid of visual, thermal and strain sensors to continuously monitor not just the end products being produced but also the health of the manufacturing machine. Such a networked environment is vital for sustainable quality control and timely and predictive maintenance of the manufacturing equipment. Such a network also ensures continuous connectivity of people in all levels of a factory, from the shop floor to upper management, [3] in the day to day activities of the plant.The successful implementation of CEM systems in a factory requires the location of sensors in and around the equipment, and open-ended software architecture to utilize the information provided by the networked sensors.

A significant amount of research has been conducted into the development of architectures for the command and control of the sensors [ [2], [4]]. Locating of the sensors is performed by mechanical fastening or adhesion and in other cases by embedding the sensors into the body of the object in question. These sensors are then monitored using wireless sensing equipment [5] to obtain information on structural health, thermal signatures, stress and strain measurements, displacement, and other performance critical341data [6].

The focus of this paper is to bring to light the use of additive manufacturing methods to embed sensors directly into the body of a part without adversely affecting its function and cost.The use of CEM leads to smarter, more connected processes, for 1) agile and efficient production; 2) manufacturing robotics that work safely with people in shared spaces; 3) computer-guided printing or casting of composites and; 4) actively controlled buildings and structures to improve safety by avoiding or mitigating accidents.1.2.Additive Manufacturing:Additive manufacturing is the process of building complex three dimensional parts from their CAD models rapidly in a layer by layer manner. CAD models are sliced to obtain tool path and layer information which are used by the additive manufacturing equipment [7]. Material is deposited in a layer form one over the other and fused to the previous layer to form a cohesive bond.

Material is deposited in a layer form one over the other and fused to the previous layer to form a cohesive bond. In the past, rapid prototyping technologies were primarily used to make parts to check for form, fit, and aesthetics but, with limited functional capabilities [8]. However in the recent past there has been a push to use these prototyping technologies to make end products with more functional applications in mind. Many researchers have successfully shown that additive manufacturing technologies can be used to manufacture end products and have termed them rapid manufacture and rapid tooling [ [9], [10], [11]]. Rapid manufacturing exhibits significant advantages over conventional manufacturing processes includingReduced lead times  Zero tooling costs  Design and redesign flexibility  Material flexibility  Part complexityThe aim of this paper is to study the advantages of using additive manufacturing to develop CEM systems.

These advantages enable their use in a multitude of direct manufacturing applications to produce parts used in aerospace, medicine, research, defense, and other sectors [12]. It is believed that many of the component items in a modern sonar transducer can be produced using these methods. The roles of additive manufacturing technologies include1. Embedded sensors 2. Print labels/EPCs 3. High performance temperature sensors 4. Fabricate on demand 5. Part DNA 6. Designing and producing conforming transducers 7. Tissue, organ production and other medical applications 8. Manufacture of assembled actuated mechanism 9.

Fabricate on demand 5. Part DNA 6. Designing and producing conforming transducers 7. Tissue, organ production and other medical applications 8. Manufacture of assembled actuated mechanism 9. Manufacture of micro – sensors and micro - actuatorsAdditive manufacturing in CEM can also lead to sustainable mass production of smart fabrics and other wearables with applications in many areas, such as electronics to provide versatility without recourse to a silicon foundry, and the processing of emerging materials such as carbon fiber and polymers that offer the potential to combine capability for electrical and342optical functionality with important physical properties including strength, durability, and disposability. In the forthcoming sections we will discuss the different rapid prototyping processes and their ability to aid CEM.2. Discussion: The roles of Additive Manufacturing in CEMAs mentioned before additive manufacturing can be a boon to promote CEM.

Discussion: The roles of Additive Manufacturing in CEMAs mentioned before additive manufacturing can be a boon to promote CEM. We will now discuss in detail the capabilities of additive manufacturing to aid CEM and their use in various applications and we will corroborate the same using research being conducted into the respective areas.2.1.Embed sensorsOne of the most critical aspects of Cyber Enabled Manufacturing is the embedding of wireless sensors into parts to accurately monitor changes in temperature and stress that the component is subjected to. Research conducted has shown that performance of wireless sensors does not deteriorate when embedded in materials [5]. Other types of sensors used include Fiber Bragg Grating, (FBG) which has shown promise for use in these applications [13]. The embedding is not restricted to the type and size of the sensors, and multiple additive manufacturing processes are available for this purpose, as discussed here.

The embedding is not restricted to the type and size of the sensors, and multiple additive manufacturing processes are available for this purpose, as discussed here. In some cases the sensors are directly inserted into the part, while in others, place holders are inserted that can be removed to include sensors after part manufacture.2.1.1. Selective Laser Sintering (SLS):Research conducted by Maier et al. has shown that selective laser sintering can be used to embed FBG strain gauges into sintered parts thereby encouraging CEM and remote sensing. Figure 1 and Figure 2 are used to explain the research.

has shown that selective laser sintering can be used to embed FBG strain gauges into sintered parts thereby encouraging CEM and remote sensing. Figure 1 and Figure 2 are used to explain the research. An ISO 3167 tensile test specimen was made with an embedded FBG sensor to measure strain change.Figure 1: Schematic of an ISO 3167 specimen with the location of the installed fiber and FBG [14]343Figure 2: Top and side view of powder bed with component and fiber location [14]The fiber and the component under construction are located on the powder bed using place holders as shown in Figure 2. The fibers are protected from the sintering process by the use of polyimide jackets to withstand the temperatures (220oC) produced during the build. The research has led to successful construction of fiber embedded tensile test specimens that are stable, and creep free, using SLS.2.1.2.

The research has led to successful construction of fiber embedded tensile test specimens that are stable, and creep free, using SLS.2.1.2. Shape Deposition Manufacturing (SDM):Shape deposition manufacturing is another additive manufacturing process that is showing applications in embedding sensors into parts. Cham et al. [15] conducted a feasibility study of using SDM to embed components and sensors into parts produced by this additive manufacturing method. Their conclusion was that SDM can be used to make mechatronics components with embedded sensors that would be difficult to make using conventional manufacturing processes.

Their conclusion was that SDM can be used to make mechatronics components with embedded sensors that would be difficult to make using conventional manufacturing processes. Their research is corroborated by work done [ [16], [17]] at Stanford University into embedding fiber optic and thin film sensors into components at critical locations and measuring the sensory output.344Figure 3: Procedure for embedding thermo-mechanical thin film sensors [16]Figure 3 shows the procedure for embedding a thin film thermo–mechanical sensor into metallic components. One of the advantages of using SDM is its usability with metallic materials [18]. Thin film sensors were characterized using four-point bend tests which compared favorably with commercially available strain gauges and theoretical models.Figure 4: Embedded fiber optic sensors [16]Figure 4 shows the embedding of optical sensors into metal matrices.

They were used toaccurately measure temperature and strain values measured on their parent components.2.1.3. Ultrasonic Consolidation (UC):Ultrasonic consolidation is a more recent addition to the additive manufacturing list. It creates metallurgical bonds between its layers through the use of ultrasonic energy to weld thin sheets of metal onto the preceding material. It is coupled to a CNC mill that removes and shapes excess material as and when necessary. This leads to the production of highly dense parts that345can be made of metal. Siggard et al. [ [19]] have successfully demonstrated the use of UC to embed sensors into dense metallic parts.Figure 5: Sensor embedded using alumina filled potting epoxy [19]Figure 6: Ultrasonic consolidation of aluminum onto embedded sensor [19]They use alumina filled potting epoxy to embed sensors into pre-milled slots in metallic parts as shown in Figure 5.

Any excess material is milled off to produce a flat surface onto which thin sheets of metal are ultrasonically welded as shown in Figure 6. Room temperature and high temperature 300oF embedding was studied with the high temperature embedding exhibiting favorable results without any loss in the sensor’s performance characteristics.2.2.Direct manufacture of assembled actuated mechanismsConventional manufacturing processes are incapable of creating linkages, making additive manufacturing an important tool in the quick creation of mechatronics assemblies that can be actuated.2.2.1. PolyJet 3D PrintingUnlike conventional manufacturing processes, some additive manufacturing processes provide the flexibility to use multiple materials for part generation using the same process during a single run.

One such process is the PolyJet 3D Printing process that contains multiple jets, each depositing different materials, to generate integrated assemblies featuring stiff components and flexible joints and gaskets. Researchers at Virginia Polytechnic Institute and State University have taken this a step further to include fibers to the existing process to create jointed and actuated mechanisms [20].Figure 7: Actuated finger created using Objet PolyJet 3D Printing Machine [20]346The build process is halted to include the fibers and then continued to create an assembly of a finger that has both rigidity and dexterity without the need to change processes. The fibers are placed into pre-designed channels and the build process resumes creating an assembly capable of actuation as shown in Figure 7.2.2.2. Stereo Lithography (SL)Stereo Lithography is the process of curing photo-curable polymers contained in vats in a layer by layer manner to form 3D shapes.

Stereo Lithography (SL)Stereo Lithography is the process of curing photo-curable polymers contained in vats in a layer by layer manner to form 3D shapes. Significant research has been conducted into SL for its use to embed sensors and actuators. In this case, inert inserts are placed into the vat of photo- polymer before or during the build of the 3D part [21]. These inserts are later removed to be replaced by the sensors or actuators that were pre-determined to be placed. Work conducted by Tse et al. indicates its application includes micro-sensor packaging [22]. De Laurentis et al. created complex working mechanisms in a single run SL process without the need for external assembly. Experimentally determined clearance values, and support structure generation, were used to make mechanical assemblies.

Experimentally determined clearance values, and support structure generation, were used to make mechanical assemblies. Figure 8 and Figure 9 show the results they produced.Figure 8: Universal joint built using SL [23]Figure 9: Robotic thumb built using SL [23]2.3.Manufacture of micro sensors and actuatorsThe need of the hour is the manufacture of sensors that do not adversely affect the functioning of the entity that they are mounted on. If the act of sensing affects the applicability347and basic functions of the part, it is detrimental to the system the part is used on and increases cost. Therefore the need to manufacture sensors that are small enough to be included in the component is great. Additive manufacturing plays a vital role in enabling the direct manufacture of micro–sensors and micro–actuators. Micro-Stereo lithography (MSL) is an additive manufacturing process that enables the fabrication of sensors of small size.2.3.1.

Micro-Stereo lithography (MSL) is an additive manufacturing process that enables the fabrication of sensors of small size.2.3.1. Micro–StereoLithography (MSL):Research conducted by Leigh et al. has highlighted the use of MSL in the fabrication of flow sensors of small dimensions. Since resins used in MSL are required to provide more functionality than regular structural strength they have used an MSL resin, embedded with magnetite (Fe3O4) nanoparticles, to fabricate a flow sensor device. The device fabricated is shown in Figure 10.Figure 10: Flow sensor device fabricated using MSL [24]The advantages of MSL include small size and easy dimensional and functional adaptability to suit situational needs. The electronics for the device are not altered in the case of a redesign and hence the costs incurred during redesign are minimal. Additive manufacturing techniques like MSL enable the rapid and cost effective manufacture of such small devices even on a small scale.

Additive manufacturing techniques like MSL enable the rapid and cost effective manufacture of such small devices even on a small scale. Research conducted in South Korea [25] furthers the use of MSL in the manufacture of microscopic bellow and micro–grippers. These micro-actuators can be used in various applications including biotechnology, biochemistry, and micro-sensing. Figure 11 shows microscopic images of the bellows manufactured using the MSL process.348Figure 11: Microscope images of micro bellows developed using MSL [25]To further test this capability they manufactured two–grip micro-grippers shown in Figure 12, and three–grip micro-grippers shown in Figure 13.

The micro-grippers were successfully tested to show their bendability.Figure 12: Fabrication and testing of two-grip micro-gripper [25]Figure 13: Fabrication and testing of three-grip micro-gripper [25]2.4.Printed circuits/Embedded Passive Component’s (EPC)In recent times it has been widely acknowledged that the use of EPC’s in electronics produces a 30% reduction in space requirements, better signal integrity and significant cost savings over discrete surface mount devices. For the manufacture of EPC’s, use of additive manufacturing provides much higher flexibility with respect to material composition and layer349thickness as compared to traditional manufacturing processes [26]. Inkjet printing is thought of to be a promising additive manufacturing technology in this regard.2.4.1.

Inkjet printing is thought of to be a promising additive manufacturing technology in this regard.2.4.1. Inkjet printingIrrespective of the manufacturing process involved, tolerances of embedded resistors after manufacturing exceed 15%, bringing into the equation post process laser trimming which, in turn, increases costs. Embedded sensors based on polyimide were developed by Dupont as ink material. With the use of ink jet printing the tolerance variation of the embedded sensors was expected to be reduced to 1%, thereby eliminating the need for laser trimming. This research was further corroborated by work done at Nokia [27]. In comparison to etching and thin-film technologies, ink jet printing offers the added advantage of using multiple materials on different print heads during the same cycle reducing the number of processes required during the printing of EPC’s.

The authors, while acknowledging the advantages of using inkjet printing, have used six sigma methodologies to further improve the process and making it more attractive for commercialization.2.5.In vivo surgery2.5.1. Bio Laser Printing (BioLP)Computer Assisted Medical Intervention (CAMI) is the process of providing tools and data to the clinician to aid in the performance of diagnostic and therapeutic actions like surgery and local injection of drugs in an accurate and safe manner. Once perfected, this system will be a major advancement in the field of medicine and will lead to increased use of robots and automation in surgery. As proof of concept, researchers in France [28]have successfully performed the first in vivo bio-printing in mice. BioLP as an additive manufacturing process was chosen to aid in bone repair. BioLP, like other additive manufacturing processes, has rapid production rates and the ability to easily customize based on immediate need.

BioLP, like other additive manufacturing processes, has rapid production rates and the ability to easily customize based on immediate need. Figure 14 shows the equipment and setup for in vivo bio printing.Figure 14: a) Schematic of setup for in vivo laser printing b) Specific holder for in vivo printing in mouse calvaria defects [28]350The researchers successfully completed the surgery and post process MRI scans revealed good tissue regeneration and lack of any necrosis. Although the application was only in bone surgery, the success achieved is significant in the implementation of CAMI.2.6.Organ printing and labelingOne of the major challenges in today’s world of medicine is human casualty associated with organ failure. The reasons for this are innumerous including lack of organs available for transplant, time constraints involved in organ processing, incompatibility between donor and patient, etc.

The reasons for this are innumerous including lack of organs available for transplant, time constraints involved in organ processing, incompatibility between donor and patient, etc. The advent of additive manufacturing has led to its widespread use in medicine, especially in the areas of organ printing [ [29], [30]]. But a system to accurately monitor the condition of the organs and ensure their timely and safe usage is as yet unavailable. Researchers at the University of Nebraska–Lincoln have proposed the use of RFID tags embedded in the organs to monitor their vital signs and guard against tissue necrosis [31]. Additive manufacturing systems are currently available in the market for bio-printing and for Radio Frequency Identification (RFID) printing.

Additive manufacturing systems are currently available in the market for bio-printing and for Radio Frequency Identification (RFID) printing. It is theorized that a hybrid system could be designed and built that could perform both the task of printing the organ and also non-invasively printing the RFID tag.2.7.Ultrasound transducersTransducers are components that convert electrical energy into ultrasound and vice versa. They are used widely in ultrasound imaging devices but are, as yet, difficult to manufacture, owing to the intricate nature of the design and the ceramic materials used in construction. Additive manufacturing in the form of digital micro printing provides an easier solution to the manufacture of transducers. Researchers at General Electric [ [32], [33]] have successfully created transducer components out of additive manufacturing and are working towards improving the process to make it even more cost effective.3.

ConclusionThis paper is aimed at bringing to light the use of additive manufacturing in enabling cyber enabled manufacturing. Various additive manufacturing processes were investigated and their applications in the fields of sensor embedding, micro-sensor manufacturing, transducer manufacturing, actuator and assembly manufacturing, micro-actuator manufacturing and computer assisted medical intervention were studied. It is found that additive manufacturing, owing to its advantages of low manufacture time, easy customizability, material flexibility, and greater tolerance of part design complexity, is a viable option for use in cyber physical systems to bring about cyber enabled manufacturing.4. References[1] R. Rajkumar, I. Lee, L. Sha and J. Stankovic, "Cyber - Physical Systems: The NextComputing Revolution," in Design Automation Conference, Anaheim, California, 2010. [2] L. Wang, P. Orban, A. Cunningham and S.

Lee, L. Sha and J. Stankovic, "Cyber - Physical Systems: The NextComputing Revolution," in Design Automation Conference, Anaheim, California, 2010. [2] L. Wang, P. Orban, A. Cunningham and S. Lang, "Remote real-time CNC machining forweb based manufacturing," Robotics and Computer-Integrated Manufacturing, vol. 20, pp. 563-571, 2004.[3] M. Takano, "Human-Centered Industrial Control Systems through Seamless351Communication and Flexible Reconfiguration," in SICE Annual Conference, Japan, 2008.[4] Q. Hao, W. Shen and L. Wang, "Towards a cooperative distributed manufacturing management framework," Computers in Industry, vol. 56, pp. 71-84, 2005.[5] J. P. Lynch and K. J. Loh, "A Summary Review of Wireless Sensors and Sensor Networksfor Structural Health Monitoring," The Shock and Vibration Digest, vol. 38, no. 2, pp. 91- 128, 2006.[6] V. Giurgiutiu and A.

Hopkinson and P. Dickens, "Rapid protoyping for direct manufacture," RapidPrototyping Journal, vol. 7, no. 4, pp. 197-202, 2001.[10] Y. L. Hou, T. T. Zhao, C. H. Li and Y. C. Ding, "The Manufacture of Rapid Tooling by Stereo lithography," Advanced Materials Research, Vols. 102-104, pp. 578-582, 2010. [11] L. Hao and S. Dadbakhsh, "Materials and Process Aspects of Selective Laser Melting ofMetals and Metal Matrix Composites: A Review," Chinese Journal of Lasers, vol. 36, no. 12, December 2009.[12] V. Petrovic, J. V. H. Gonzalez, O. J. Ferranda, J. D. Gordillo, J. R. B. Puchades and L. P. Grinan, "Additive layered manufacturing: sectors of industrial application shown through case studies," Internation Journal of Production Research, vol. 49, no. 4, pp. 1061-1079, 2011.[13] Y. Rao, "Study on Fiber-Optic Low Coherence Interferometric and Fiber Bragg GratingSensors," Photonic Sensors, vol. 1, no. 4, pp. 382-400, 2011.[14] R. R. J. Maier, W. N. MacPherson, J. S. Barton, M.

2, p. 8, 2010.[29] F. P. W. Melchels, M. A. N. Domingos, T. J. Klein, J. Malda, P. J. Bartolo and D. W.Hutmacher, "Additive manufacturing of tissues and organs," Progress in Polymer Science, vol. 37, pp. 1079-1104, 2012.[30] J. Giannatsis and V. Dedoussis, "Additive fabrication technologies applied to medicine and health care: a review," International Journal Advanced Manufacturing Technology, vol. 40, pp. 116-127, 2009.[31] E. C. Jones, M. Cavitt and D. Kong, "Manufacturing Feasibility Evaluation of RFID chipsembedded in aritficial organs," in International Conference on Future Information Technology and Management Engineering, 2010.[32] "http://ge.geglobalresearch.com/blog/additive-manufacturing-building-blocks-for-the-future/," [Online]. [Accessed 10 July 2012].[33] "http://www.economist.com/node/18710080," [Online]. [Accessed 8 July 2012].353.

plications of Powder Densification Maps to Direct Metal SLS/HIP ProcessingMartin Wohlert†, David L. Bourell†*, Suman Das*, Joseph J. Beaman†*†Texas Materials Institute Laboratory for Freeform Fabrication *Mechanical Engineering Department University of Texas at Austin Austin, TX 78712Abstract Recent investigations have shown that Direct Selective Laser Sintering in conjunction with Hot Isostatic Pressing (HIP) is capable of producing fully dense, near net shape, high performance metal components. A key aspect of this combined processing method is the selection of appropriate HIP process parameters to obtain full density parts. Powder Densification Maps provide a powerful tool for predicting the effect of time, temperature, pressure, and material properties on part density. This paper summarizes developments in Powder Densification Map production for Ti-6Al-4V and Inconel 625 materials.

This paper summarizes developments in Powder Densification Map production for Ti-6Al-4V and Inconel 625 materials. A comparison of theoretically predicted and experimentally determined densities for a variety of processing conditions is presented.IntroductionThe material densification rate during Hot Isostatic Pressing (HIP) is dependent on a variety of factors including: intrinsic material properties such as yield strength and diffusion rates; sample specific variables including the initial density, part geometry, and powder particle size; and HIP processing parameters such as temperature and pressure. The ability to model the densification process is beneficial and permits planning of more advanced process optimization such as control and modification of microstructure or geometric distortion.The influence and interaction of the varying mechanisms exhibited across the entire processing envelope is most readily represented in a graphical format.

This graphical representation has become known as a HIP map. The basis of this concept was developed by Weertman [1]; however, the construction of the maps was brought to fruition by Ashby [2] in the early 1970’s. The constitutive equations for the generation of the maps have been incorporated into a computer program which was used to create the deformation maps provided here. In the following sections, the deformation mechanisms incorporated into the model and the accompanying constitutive equations will be reviewed. For two materials selected for SLS+HIP trials (Ti-6Al-4V and Alloy 625), material property data and calculated maps will be shown. Experimental data obtained from glass encapsulated HIP trials will be presented to validate the150HIP map results.

Experimental data obtained from glass encapsulated HIP trials will be presented to validate the150HIP map results. A discussion of the accuracy of the theoretical predictions relative to the experimental results will be presented.Powder DensificationThe models developed for both conventional pressureless sintering and HIP assume an initial structure consisting of uniformly sized, perfectly spherical particles. This assumption greatly simplifies the calculations in comparison to a system of arbitrary, irregularly shaped particles, and in fact, is not too far removed from the powders selected for SLS+HIP. Initially the spherical particles are assumed to be in point contact as shown in Figure 1.

Initially the spherical particles are assumed to be in point contact as shown in Figure 1. As the sintering or HIP process proceeds, the contact area between particles increases to form “necks” or “bridges” between adjacent particles.Initial point contact (Spherical particles)Early neck formation (Grain boundary formed between particles)Late stage neck growthFinal stage (Fully coalesced spheres)Figure 1. Model of sintering process for uniformly-sized, spherical particles (after German) [3]In early stages of sintering, particle necks provide strength, but very little densification takes place. As sintering proceeds, necks continue to enlarge, and particles begin to coalesce, resulting in increasing bulk density. The enlarged necks between particles produce a cylindrical pore structure throughout the powder mass typified in Figure 2a.

The enlarged necks between particles produce a cylindrical pore structure throughout the powder mass typified in Figure 2a. While this interconnected pore structure exists, it serves as an important conduit for mass transport; however, when a theoretical density of approximately 92% is reached, the continuous, cylindrical structure becomes unstable and collapses to form isolated spherical pores, as illustrated by black dots in Figure 2b. This transition is defined as the boundary between the initial sintering conditions, Stage I, and the later Stage II process. This change in sintering behavior requires a modification of the equations used to described the densification process.

This change in sintering behavior requires a modification of the equations used to described the densification process. As a result, two different equations exist for each densification mechanism.151(a) Figure 2(b)Illustration of the (a) cylindrical Stage I pore structure and (b) isolated, spherical Stage II pore structure [4]Driving Force for DensificationFor pressureless sintering, a reduction in surface area is the only driving force for the densification of the powder mass. In the case of HIP, an additional driving force stemming from the stress created by the compaction pressure is also active. To provide a convenient and unified form of the two driving forces, the surface area component of the driving force is converted to an effective stress and summed with the net applied HIP gas pressure, as shown in Equations 1 and 2.

With typical HIP pressures on the order of 100 MPa, it is apparent that there is a substantial driving force increase associated with HIP compaction relative to conventional pressureless sintering.Densification Mechanisms and Map ConstructionIn all stages of sintering, several densification mechanisms act simultaneously. In Ashby’s initial analysis for pressure sintering maps, he included four independent mechanisms: lattice diffusion, grain boundary diffusion, power-law creep (or dislocation creep), and plasticity (yielding/dislocation glide). In more recent work, the additional mechanisms of Nabarro-Herring creep and Coble creep were included for cases where the internal grain size of the particles is substantially smaller than the particle diameter [6].

Grain growth and grain boundary separation from pores during Stage II densification are also incorporated into the model.Having successfully defined constitutive equations for the densification mechanisms, calculation of the HIP maps may proceed. A flow chart of the process is shown in Figure 3. To begin the process, the values for all required material properties must be entered. Next, a starting temperature, pressure, and initial relative density D o are set. The program calculates the densification rate contribution of each deformation mechanism and sums them to obtain an overall densification rate D & total. Then, an incremental increase in the relative density D is / D & total. selected and the time (D The dominant densification mechanism, the new relative density and the total elapsed time are recorded, and the entire process is repeated starting at the newly incremented density. Eventually, the relative density is incremented to the full theoretical density.

Eventually, the relative density is incremented to the full theoretical density. At this point, a new temperature is selected, the relative density is reset to its initial value, and the entire process is repeated. The collected data is then used to generate a plot of relative density (D ) versus temperature (T) for a given time in the form of isochronal lines. The information collected during each iteration on the dominant deformation mechanism may be indicated in appropriate regions on the map. A complete listing of the data used in constructing the HIP maps is given in the Table.t) required to obtain the densification is calculated from Dt=DExperimental Verification of the HIP ModelThe densification predictions of the HIP model were verified through a series of HIP experiments. Alloy 625 and Ti-6Al-4V powders were processed over a range of temperatures, pressures, and times.

Alloy 625 and Ti-6Al-4V powders were processed over a range of temperatures, pressures, and times. The density achieved for a particular set of processing conditions was153StartPext=0?NoCalculate Dreplace DYesDetermine Stage I or Stage IIDetermine if pore separation has occurredIntegrate the appropriate V b equation to get a new G or RCompute the Driving PressureCalculate all D • to obtain D(volume, boundary, Coble, P-L, N-H) and sum total and the dominant deformation mechansismAssume a Dand calculate a new timeNoIs D =1?YesNoAnother Temp.

dependence of Y.S.0.02 MPa/K0.82 MPa/KAtomic Volume1.09x10-29 m31.76x10-29 m3Power Law Creep Exponentn82Power Law Creep Activation EnergyQc269 kJ/mol225Power Law Creep Reference Stressref900 MPa100 M PaVolume Diffusion Pre-exponentialDV1.2x10-3 m2/s6.6x10-9 m2/sVolume Diffusion Activation EnergyQV259 kJ/mol169 kJ/molBoundary Diffusion Pre-exponentialDB7.17x10-3 m2/s1.2x10-5 m2/sBoundary Diffusion Activation EnergyQB209 kJ/mol97 kJ/molGrain Boundary Width6x10-10 m6x10-10 mmeasured by gas pycnometry or Archimedes’ principle and compared to the value predicted by the HIP maps. Powder was placed in borosilicate glass, evacuated to at least 5x10-5 Torr, degassed at 350(cid:176) C for a minimum of 24 hours to remove the majority of surface adsorbed gases as well as some species dissolved within the powder [7], and the glass tube was sealed prior to HIP.

The glass capsules were processed in an ABB QIH-3 HIP system equipped with a graphite heating element. The process cycle began with a 10(cid:176) C/min ramp from room temperature to the155desired processing temperature. Once the operating temperature was reached, the pressure within the HIP vessel was raised. After the desired processing time had elapsed, the temperature was decreased at 15(cid:176) C/min and the pressure was gradually reduced to atmospheric pressure.Results and DiscussionA HIP map was generated for the Alloy 625 material as shown in Figure 4. All three operating parameters (time, temperature, and pressure) were varied in the HIP experiments. In most cases, the predicted values are within one or two percent of the experimentally determined density.1B.D. (II)Volume Diff. (II)0.95N-H Creep (II)0.9Volume Diff. (I)Boundary Diff.

(II)Volume Diff. (II)0.95N-H Creep (II)0.9Volume Diff. (I)Boundary Diff. (I)0.85y t i s n e D e v i t a l e R0.80.75Power Law Creep (I)0.71 hour8 hours0.65Yielding0.60 0 70 5 70 0 80 5 80 0 90 5 90 0 0 10 5 0 10 0 1 1Temperature, °CFigure 4HIP map for Alloy 625. HIP pressure = 10 ksi. Experimentally determined values have been plotted for 1 hour (l) and 8 hours (n).156A similar set of experiments and HIP maps were generated for the Ti-6Al-4V material. The results are shown in Figure 5. The short process cycle data (5 min) still show significant deviation from the theoretical values. It has been clearly demonstrated by a number of researchers that fine grained (<10 m m) Ti-6Al-4V exhibits superplastic behavior at temperatures near 900(cid:176) C [8,9]. These authors have suggested that a grain boundary sliding (GBS) mechanism accounts for the behavior of the material under the specified processing conditions.

These authors have suggested that a grain boundary sliding (GBS) mechanism accounts for the behavior of the material under the specified processing conditions. This mechanism is currently not included in the HIP Map model. A number of constitutive relationships have been developed to describe deformation associated with grain boundary sliding (or shearing). The incorporation of an appropriate rate equation into the model could improve the prediction accuracy for superplastic materials such as Ti-6Al-4V.1Boundary Diffusion (II)0.95Coble Creep (II)PLC (II)0.90.85y t i s n e D e v i t a l e R0.80.75Boundary Diffusion (I)Power Law Creep (I)0.70.650.1 hourYielding2 hours0.64 hours0 0 50 5 50 0 60 5 60 0 70 5 70 0 80 5 80 0 90 5 90 0 0 1Temperature, °CFigure 5HIP Map for Ti-6Al-4V. HIP pressure = 10 ksi.

HIP pressure = 10 ksi. Experimentally determined values have been plotted for 0.1 hour (l), 2 hours (s), and 4 hours (n).157Summary and ConclusionsHIP maps were generated to predict the densification behavior of Alloy 625 and Ti-6Al-4V. The ability to accurately determine an appropriate HIP cycle without substantial trial-and-error testing could greatly reduce the time and cost associated with the development of SLS+HIP processes for new materials. The HIP maps were constructed by calculating the densification contributions of several independent deformation mechanisms: lattice diffusion, boundary diffusion, power-law creep, Nabarro-Herring creep, and Coble creep. The density predictions of the HIP map model were verified by experimental HIP cycles run on glass encapsulated powders.In general, the results of the Alloy 625 and Ti-6Al-4V modeling are quite illustrative of the benefits and limitations of the mechanistic model approach.

As was demonstrated, the model typically provides a useful qualitative description of the densification rate without requiring a high degree of accuracy in the material property data. The quantitative results, however, are highly dependent on the material property data selected for the model. For most materials it is found that power law creep dominates the densification process over a wide range of processing conditions. Boundary diffusion provides a substantial supporting role, particularly at moderate temperatures. This suggests that densification rates at moderate processing temperatures may be highly dependent on particle size. Additional experiments, measuring the effect of particle size variations on densification rate could be useful to further verify the accuracy of the HIP maps.References[1] Weertman, Trans Met. Soc AIME. 227,1475 (1963). [2] M. F. Ashby, “A first Report on Deformation-Mechanism Maps,” Acta Metallurgica, Vol.20 (1972) pp. 887-897.[3] Randall M.

Research and applications conducted over the past few years are described including: Helisys Laminated Object Manufacturing System (LOMS) fabrication of: ceramic piezoelectric actuators, tooling for multifunctional materials, and anatomical prototypes for surgical visualization; fabrication of mesoscale electronic and sensor components using a laser forward transfer direct write technique; and visualization of complex, 3-D microstructures using a Stratasys Fused-Deposition Modeler. The paper closes with a brief overview of future SFF related work at the NRL.IntroductionSolid Freeform Fabrication (SFF) and related techniques are used at the Naval Research Laboratory (NRL) for a variety of materials related investigations.

This paper will describe several recent SFF projects conducted using an NRL Helisys Laminated Object Manufacturing System (LOMS Model 2030H) or a custom NRL built laser-based, direct material writing system known as: “Matrix-Assisted-Pulsed-Laser-Evaporation Direct-Write” (MAPLE DW). Use has also been made of the Stratasys Fused-Deposition Modeler (FDM) in some collaborative efforts. We begin with a description of several LOMS-based applications including: ceramic piezoelectric actuator fabrication, rapid tooling for multifunctional unmanned air-vehicle components, and anatomical prototyping for surgical visualization. Fabrication of mesoscale electronic and sensor components using the MAPLE DW system is described next followed by a description of the use of FDM for creating a physical model for complex, 3-D material microstructures.

We conclude with a brief overview of future SFF related work at the NRL.Laminated Object Manufacturing System ApplicationsPiezoelectric ActuatorsAdvancing the development of high authority (i.e., high force and displacement) piezoelectric actuators through the use of innovative designs and advanced fabrication techniques is a critical enabling step for military systems [Wu et al., 1999]. The NRL has been involved in a program to develop the Navy’s SFF capability for rapid R&D of novel tailored actuators as a means of accelerating their insertion into fleet applications.One such novel actuator, the telescoping actuator, patented by Lewis and Kahn [1997] (Figure 1), serves as an active element in machinery mounts for noise (acoustic signature) reduction. The telescoping design provides for displacement amplification with an output force proportional to the cross-section of the smallest element.

The telescoping design provides for displacement amplification with an output force proportional to the cross-section of the smallest element. These actuators are used in constrained spaces of varying size and have varying power/displacement requirements. Another actuator,576designed for active payload vibration suppression during spacecraft launch, consists of a thin ceramic piezoelectric ring with small axial holes along the circumference for weight savings.Actuators can be fabricated using an injection-molding process, but this is too expensive and time consuming for multiple design evaluations. Instead, a modified LOMS process has been used to rapidly fabricate functional components for direct evaluative testing. Griffin et al.

Instead, a modified LOMS process has been used to rapidly fabricate functional components for direct evaluative testing. Griffin et al. [1994] document some of the early efforts to adapt the LOMS technique for fabricating ceramic (alumina) components.The telescoping actuators are built by integrating the lamination of thin lead-zirconate- titanate (PZT) embedded polymer sheets [Bender et al., 2000] with precise laser cutting by the LOMS. Engineering of the ceramic-polymer sheet stock, referred to as ceramic tape, is critical to the success of the process. The ideal ceramic tape is stiff enough to avoid excessive distortion during handling, amenable to laser cutting, and suitable for lamination via solvent welding, while possessing appropriate firing properties [Cawley and Liu, 1998].PZT tapes formulated with a polymer binder system optimized for alumina tapes were tried first.

These tapes were very flexible due to the large amount of plasticizer that was added to allow mixing with the high specific gravity PZT. Microstructural characterization of the tape using scanning electron microscopy (SEM) showed non-uniform distributions of binder and PZT with PZT-free regions and a vertical density gradient (porous to dense going from top to bottom). Actuators made with these tapes were badly cracked and deformed (Figure 2A). The cracking and deformation were greatly reduced (Figure 2B) using tapes with PZT-optimized binder, which facilitates more uniform PZT powder and binder distributions within the tape.Selection of an appropriate solvent for lamination welding is vital to achieving a good bond between the tape layers. The solvent must act as a tackifier, enabling bonding between the layers, and it must have a relatively low vapor pressure so that it doesn’t evaporate before affecting the lamination process.

The solvent must act as a tackifier, enabling bonding between the layers, and it must have a relatively low vapor pressure so that it doesn’t evaporate before affecting the lamination process. Also, its action must be limited to the interface region. Otherwise it can diffuse to the laser-cuts and cause partial rebonding, which makes decubing of unwanted material from the build more difficult.Previous research indicated that a light spray of propanol led to satisfactory lamination [Bender et al., 2000]. On closer examination, however, small interlaminar debonds were observed scattered throughout the actuator. One reason for the presence of these defects was uneven evaporation of the propanol. Longer spray times were tried in an attempt to prevent uneven distribution of the propanol, but the presence of too much solvent created small pockets of liquid that could not be eliminated during lamination leading to even larger interlaminar defects.

A low vapor pressure solvent, iso-amyl alcohol, was tried as a laminating solvent. Uniform distribution of the solvent was observed, and after firing, fewer defects were observed. However, the iso-amyl alcohol softened the ceramic tape binder too much making decubing difficult. A solution was found that required adding 10% polyethylene glycol (PEG- MW 200) to the propanol to retard the evaporation, leaving behind a thin tacky interlaminar film. SEM characterization revealed a continuous interlaminar film ~1 µm in width was left behind with no evidence of interlaminar defects. The PEG has time to dissipate away from the interfaces during binder burnout leading to low defect actuators after firing (Figure 2B).Understanding the important SFF process parameters related to the PZT tapes and solvent lamination has allowed us to develop adequate process control, which leads to dimensional control and property reproducibility.

As a result, this LOMS-based SFF process will facilitate more rapid R&D of novel ceramic piezoelectric actuators.577Rapid Tooling for Multifunctional Materials FabricationThe LOMS machine has also been used to make molds and patterns for use in fabricating multifunctional fuselage prototypes for a micro air vehicle (MAV). We are integrating the MAV’s structure and battery functions into one material to reduce weight and increase flight- time [Thomas et al., 2001]. The novel structure-battery materials being developed for this application are comprised of a thin-layer, polymeric lithium-ion battery material with structural additives and a barrier-layer packaging1. Molds and patterns are needed to form the composite structure-battery material “package” into the complex, 3-D curved fuselage shell shape.A LOMS pattern of the MAV fuselage and the corresponding molds for the upper and lower fuselage halves are shown in Figure 3.

These “tooling” components are made from standard Helisys paper (LPH 042). After decubing, they are sanded smooth and then epoxy coated for stiffening and strengthening at structure-battery fabrication temperatures ( max). The tooling will be used in a vacuum-bag-forming operation to mold the structure-battery materials into the MAV fuselage form. Three types of structural enhancement are being considered. The first involves adding polypropylene (PP) or polyethylene (PE) films internally and bonding with the battery material. The second will increase the thickness of the PP or PE in the barrier-layer packaging to provide the required shape stiffness. The third type uses an epoxy- fiber shell. All three concepts require MAV fuselage molds/patterns for vacuum-forming at elevated temperatures ( max) to affect interlayer bonding and shaping of the thermoplastics and curing of the thermosets.

All three concepts require MAV fuselage molds/patterns for vacuum-forming at elevated temperatures ( max) to affect interlayer bonding and shaping of the thermoplastics and curing of the thermosets. LOMS technology allows for rapid creation of complex tooling for this advanced development demonstration of multifunctional structure-power materials concepts.C°80C°80Anatomical Prototyping for Surgical VisualizationRecently, we used the LOMS to create a very detailed anatomical prototype of a human spine with severe scoliosis from CT-scan digital data (Figure 4). SFF models are routinely used in medicine today for visualization and planning of complex surgical procedures, physician training, and informing/educating the patient, and that is the intended purpose of this model. Several fractional-scale models of the spine were made at the National Naval Medical Center using their Stratasys Fused-Deposition Modeler (FDM).

Several fractional-scale models of the spine were made at the National Naval Medical Center using their Stratasys Fused-Deposition Modeler (FDM). Due to the complexity of the deformed spine structure, a full-scale model was desired, but model sizes are limited to 10 inches for this particular FDM machine. The build envelope for the NRL LOMS was large enough to fabricate the full-scale model of the spine, and it exhibited an added advantage of providing outstanding contrast on the various detailed skeletal features.××inches in dimension created with approximately 1070 layers of Helisys LPH 042 paper stacked in the 5 inch dimension. The build process took about 70 hours due to the fine-scale features in the model inch).

The build process took about 70 hours due to the fine-scale features in the model inch). Several days were spent on pre- and post- and the small crosshatching size ( build tasks and *.StL file manipulations, and decubing took approximately 30 hours spread-out over several weeks.The spine prototype, shown in Figure 4, was fabricated as a solid block1857×25.050.0The software MIMICS2 was used to create the *.StL file from the CT-scan data. The data was filtered using triangle reduction, but the resulting *.StL file was still too large at 68.8 Mb for1 The battery is ~ 0.5 mm thick in the bicell configuration [Thomas et al., 2001], and was developed in the early 1990’s by Telcordia Technologies [Gozdz & Warren, 1997]. The packaging is a multilayer polymer-aluminum heat- sealing laminate (~0.1 to 0.5mm thick) that serves as a chemical/moisture barrier layer.2 http://www.materialise.be/mimics/578the LOMSlice software.

We used Cyberware’s Decimate Polygon Reduction Software3 to reduce the *.StL file ~50% to 37.7 Mb, which was manageable by LOMSlice. The software works by coalescing polygons of “little importance” for accurate model representation via elimination of select vertices.Matrix Assisted Pulsed Laser Evaporation Direct Write (MAPLE DW)A laser-based direct material writing technique known as: “matrix assisted pulsed laser evaporation: direct write” (MAPLE DW) has been developed here at the NRL. This system, which builds on recent developments in materials and laser-based materials processing, provides a radically different method of fabricating mesoscale inorganic structures for electronic and sensor device applications and organic structures for unique biological systems applications [Piqué et al., 1999; Fitz-Gerald et al., 2000; Chrisey et al., 2000a].

An individual electronic component or an array of sensors can be built conformally over the surface of almost any substrate in air and at room temperature using MAPLE DW. This capability opens the door for the rapid prototyping and testing of customized designs.The material deposition process begins in MAPLE DW when a high-repetition-rate, 355 nm UV laser beam is focused on a material “ribbon.” Upon heating from the laser pulse, the build material transfers to the receiving substrate where it forms an adherent coating. The ribbon is a transparent tape that supports a 1-10µm thick layer of build material on one side, analogous to a conventional ink typewriter ribbon. A schematic of the process is shown in Figure 5.Using MAPLE DW, we can fabricate electronic circuit patterns with feature resolutions on the order of 10µm by synchronously moving the ribbon to a fresh, unexposed region, and the receiving substrate approximately one beam diameter.

The individual mesoscopic 3-D pixels or voxels of electronic material, one voxel transferred per laser shot, are assembled into the desired 3-D pattern. Alternating the ribbon material between a metal and a dielectric, for example, allows us to fabricate three-dimensional multilayer structures like parallel-plate capacitors.The MAPLE DW system can also be operated in a laser micromachining mode or a laser sintering mode by removing the ribbon. This provides the capability for etching grooves or vias in the substrate, or surface annealing, sintering, or etching of individual components to improve their performance or dimensional accuracy. See, for example, Pique et al. [2000] and Chrisey et al. [2000b] for improvements to the electronic properties of oxide ceramics. For infrared (IR) laser wavelengths (1 to 10µm), the penetration depth is on the order of microns for a variety of materials allowing the benefits of higher-temperature processing on thermally sensitive substrates.

Ultimately, laser annealing might allow the decomposition and reaction of organometallic precursors for transfers onto substrates that cannot tolerate processing at temperatures above the boiling point of water, (e.g., biomaterials). It might also allow some degree of bulk diffusion and sintering of ceramics for transfers onto substrates that can tolerate heating to a few hundred degrees for a short period of time (e.g., polyimide).Advances in the research community's understanding of new materials and laser-material interactions have driven the progress evident in direct writing of electronic materials. On the other hand, recent advances in the direct writing of biomaterials have been driven by progress in UV-laser based transfer technology. MAPLE DW in particular can be an extremely gentle process. Using ribbons made of an aqueous composite mixture, we have been able to transfer patterns of viable bacteria such as E.

MAPLE DW in particular can be an extremely gentle process. Using ribbons made of an aqueous composite mixture, we have been able to transfer patterns of viable bacteria such as E. coli onto various substrates [Ringeisen et al., 2001].3 http://www.cyberware.com/products/decimate.html579The capabilities offered by the MAPLE DW technique have opened up the possibility of integrating bio-structures with electronic devices on the same substrate for rapidly prototyping unique cell-based biosensors and bio-electronic interfaces. Clearly, the ability to generate components on demand, any place, on any substrate, and in a matter of minutes instead of weeks, provides us with a unique opportunity for bringing new designs to life that can only be dreamed of with today's manufacturing techniques.Fused Deposition Modeling for Microstructure VisualizationAn SFF model for the microstructure of a steel alloy is shown in Figure 6A.

This model was fabricated using the Stratasys Fused Deposition Modeling (FDM) process4. It depicts a 3-D austenite grain within the alloy where the white solid represents the cementite decorated grain boundary and cementite precipitates and the empty space represents the austenite phase. The material is a Fe-1.3%C-13%Mn model steel alloy that was isothermally heat-treated at 650°C for 50 sec. A series of 128 images of the grain was taken by optical microscopy from serial sections at 0.2µm depth increments (Figure 6B). They were read into a Solid Model Builder program developed by the University of Michigan [Marsan and Dutta, 1996]. Each image was used to generate one slice for the FDM of the solid model.The microstructures were generated and analyzed by Kral et al. [2000] at the NRL. This work was performed to develop better understanding of the true 3-D morphology and distribution of the solid-state precipitates in structural metal alloys.

[2000] at the NRL. This work was performed to develop better understanding of the true 3-D morphology and distribution of the solid-state precipitates in structural metal alloys. This understanding is important for the control of microstructure features that have a direct influence on the mechanical properties of components. Microstructural characterization via conventional microscopy provides only 2-D views from which 3-D information can be inferred, often erroneously. Three-dimensional analysis of microstructures gives accurate information on the nature of the grain boundaries, the origin of the precipitates, and the shape, dimensions and distribution of the precipitates. This information makes quantification of microstructure and its relation to material behavior more meaningful.Three-dimensional microstructures have been mathematically reconstructed using the serial sections and can be viewed in 3D using a Silicon Graphics Imaging system.

The image can be rotated and translated as it is studied. However, 3D microstructures are best studied when viewed as solid models. Microstructures inherently have complex geometries and their solid models can only be reproduced by SFF techniques thus highlighting a significant benefit of the technology.Future Work at the NRLContinued research on the MAPLE DW process and capabilities and use of the LOMS for rapid fabrication of ceramic-based components is planned. The anticipated addition of a tape- casting machine for creating ceramic-polymer precursor materials for the LOMS will lead to a number of new materials related investigations. There is also potential interest in investigating the use of SFF related techniques for creating hierarchically structured materials with enhanced and multifunctional performance.

There is also potential interest in investigating the use of SFF related techniques for creating hierarchically structured materials with enhanced and multifunctional performance. Examples might include: tailoring the fracture resistance of ceramic-polymer composites and the dielectric properties of capacitor ceramics through hierarchical patterning.4 In collaboration with the University of Michigan-Ann Arbor.580AcknowledgementsSupport for this work from DARPA, Office of Naval Research, and the Naval ResearchLaboratory Core Research Program is gratefully acknowledged.ReferencesBender, B., Rayne, R., and Wu, C., “Solid Freeform Fabrication of a Telescoping Actuator Via Laminated Object Manufacturing”, Ceram. Eng. Soc. Proc., Vol. 21, pp. 143-150, (2000).Cawley, J. D., and Liu, Z., “Applying Tape Casting to Layered Manufacturing Processes”, Ceram. Ind., Vol. 128, pp.

16th Amer. Soc. Comp., Virginia Tech, Sept. 2001, to appear.Wu, C., Lewis, D., Kahn, M., and Chase, M., “High Authority, Telescoping Actuators”, Proc. SPIE, Vol. 3674, pp. 212-219, (1999).581FiguresExpansion of the first and third cylinders plus contractionABExpansion of the third and contraction ofthe second cylinderExpansion of the first and contraction ofthe second cylinderCross Sectional View: ExcitedFigure 1: Schematic cross-sectional view of the telescoping actuator.Figure 2A and 2B: Optical micrographs showing a bottom view of the consolidated an telescoping unsatisfactory tape formulation (A) and a satisfactory tape formulation (B). Note that the outer cylinder diameter is 2.5 cm.actuators made withFigure 3. LOMS pattern and molds for a micro-air vehicle fuselage.Figure 4.

Note that the outer cylinder diameter is 2.5 cm.actuators made withFigure 3. LOMS pattern and molds for a micro-air vehicle fuselage.Figure 4. Human spine model created with the LOMS from the CT-scan data.582Alignment/ImagingLaser/LensRibbonWorkpiece Translation StageFigure 5: Schematic diagram of the MAPLE DW process.Figure 6A and 6B: (A) FDM model of cementite precipitates and austenite grain in a high carbon steel, and (B) a portion of the stack of 250 images of optical microstructures used as slices for the FDM process.583.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAPPLICATIONS OF SUPERVISED MACHINE LEARNING ALGORITHMS IN ADDITIVE MANUFACTURING: A REVIEW M. S. Joshi*, A. Flood*, T. Sparks †, and F. W. Liou *Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409 † Product Innovation and Engineering LLC, St James, MO 65559AbstractAdditive Manufacturing (AM) simplifies the fabrication of complex geometries. Its scope has rapidly expanded from the fabrication of pre-production visualization models to the manufacturing of end use parts driving the need for better part quality assurance in the additively manufactured parts. Machine learning (ML) is one of the promising techniques that can be used to achieve this goal.

Machine learning (ML) is one of the promising techniques that can be used to achieve this goal. Current research in this field includes the use of supervised and unsupervised ML algorithms for quality control and prediction of mechanical properties of AM parts. This paper explores the applications of supervised learning algorithms - Support Vector Machines and Random Forests. Support vector machines provide high accuracy in classifying the data and is used to decide whether the final parts have the desired properties. Random Forests consist of an ensemble of decision trees capable of both classification and regression.

Random Forests consist of an ensemble of decision trees capable of both classification and regression. This paper reviews the implementation of both algorithms and analyzes the research carried out on their applications in AM.IntroductionAdditive Manufacturing is a process of fabricating a component in a layer by layer manner directly from 3D CAD models by melting or sintering polymers or metal alloys using different energy sources such as laser, electron beam, arc, etc. This technology has broadened its applications in producing end-use parts because of the advantages such as the ability of fabricating complex geometries, use of diverse materials, and capacity of achieving desired mechanical properties [1].To achieve perfect structural integrity is the basic requirement in fabricating any end use part. The Generation of defects in the part during the process results in wastage of time, money, and material.

The Generation of defects in the part during the process results in wastage of time, money, and material. Additive manufacturing faces major challenges in in-situ defect detection and process control. Machine learning (ML) methods offer an opportunity to detect the defects in real time, which prevents material wastage and reduces the efforts for trial and error. Machine learning, which is seen as a subset of artificial intelligence, consists of algorithms which build a mathematical model based on sample data, known as ‘training data’, in order to make predictions or decisions on unknown ‘test data’ without being explicitly programmed to perform the task.Among different types of machine learning algorithms, supervised learning algorithms are used for classification and regression purposes. In this algorithm, the training data set contains one or more inputs and labeled desired outputs. A mathematical model is built on this data and is executed in the following steps.

In this algorithm, the training data set contains one or more inputs and labeled desired outputs. A mathematical model is built on this data and is executed in the following steps. First, the model is trained to learn a mapping function from input213to output using a training data set so that the function can predict the output for new data input. Next, the performance of the model is checked by generating predicted labels of the testing data. Last, the model is cross-validated for performance evaluation in terms of accuracy, precision, recall, etc. Output variable for classification problems is categorical, whereas regression problems have continuous real output variables [2].There are different supervised learning algorithms. This work reviews and analyzes the applications of two supervised ML algorithms- support vector machine and random forests. These algorithms were chosen as they represent the majority of research carried out using supervised learning algorithms.

These algorithms were chosen as they represent the majority of research carried out using supervised learning algorithms. Neural networks were excluded to limit the scope of this review.Supervised Machine Learning Algorithms Support Vector MachineSupport Vector Machine (SVM) is one of the powerful supervised machine learningalgorithms which can be used for classification as well as regression.In classification problems, this algorithm is used to find a decision boundary which canproperly separate unseen data into two or more categories with the help of training data.,,, ,,,,Class l• ••,,,•• ••Class 2Figure 1: Classification of linearly separable data by support vector machine algorithm [2]For linear classification of n-dimensional data into two classes, a hyperplane with (n-1) dimensions is generated. Figure 1 shows the linear classification of two-dimensional data.

Also, it is robust against the error of models and has a higher computational efficiency compared to other supervised algorithms. Due to these advantages, it has various applications in various problems such as text classification, pattern recognition, etc.Applications of SVM in Additive ManufacturingIn additive manufacturing, the SVM algorithm have been demonstrated in the following areas:1. Defect Detection:Additively manufactured components often have defects such as the incomplete fusion of the powder, porosity, cracks, inclusions, etc. These defects have a strong impact on the mechanical properties of the component. Porosity is one of the defects of major concern [3]. The existing porosity detection techniques are either visual based or simulation-based or based on post- manufacturing characterization. Vast research has been carried out on in-situ porosity detection using various sensors. Support vector machine provides great accuracy in in-situ porosity detection.

Vast research has been carried out on in-situ porosity detection using various sensors. Support vector machine provides great accuracy in in-situ porosity detection. Table 2 shows that it is used to classify the builds in two classes as defective or flawless using the input data from different sensors such as a high-speed camera, IR camera, pyrometer, etc. The following can be used as input data and class labels for training SVM model.Sr. NoInput dataModification in SVM/Kernel used Class labels1.Layer wise ImagesEnsemble classifierAnomalous/ nominal2. RGB values of images at a checkpoint Surface flatness 3.2. RGB values of images at a checkpoint Surface flatness Thermal HistoryIncreasing no. of checkpointsBad/goodLinear, Gaussian, Polynomial Kernel Normal/abnormalITable 2: Applications of SVM in defect detectionDifferent measures are taken in order to improve the performance of SVM in terms of accuracy.

In [4], A linear SVM ensemble classifier fuses visual information extracted from high- resolution layerwise images of build surface in PBF process captured from eight different sources. This classifier is trained using the labels ‘anomalous’ and ‘nominal’ automatically acquired from the post-build CT scans. An ensemble classifier combines the outputs from individual classifiers which enables multiple in-situ sensor modalities to be utilized for quality assessment. In this work, multiple images collected under different lighting conditions for each layer serve as sensor modalities to increase the accuracy of the model to 85% compared to the accuracy of 65% if individual classifiers are used.[5] Develops a methodology for DLD process based on functional principal component analysis (FPCA) to extract key characteristics from melt pool thermal images by converting them to morphological model.

Using different supervised learning algorithms, data is classified into ‘normal’ and ‘abnormal’ melt pool labels acquired via X-ray Tomography. To classify the data216Ref .[4][6][9][5]using SVM, linear, polynomial and Gaussian kernel functions are used. SVM provides the highest value of the accuracy measure (28.32%) in correctly predicting abnormal data points. Also, it is found that the polynomial kernel function provides superior performance compared to other kernels. The method carried out in this study can be applied to other processes such as PBF or EBM, which have similar energy- material interaction.Another study, [6], uses SVM to detect any possible defects occurring during the FDM process. In this study, the algorithm classifies the parts as ‘good’ or ‘bad’ using the images taken at specific check points through the automated image capturing process.

In this study, the algorithm classifies the parts as ‘good’ or ‘bad’ using the images taken at specific check points through the automated image capturing process. For training the model, section averages (of RGB values) of the images calculated at checkpoints are loaded as input to the vectors of training models. The method is capable of detecting both completion failure defects such as filament running out or printing stopped in the mid-progress and structural or geometrical defects.2. Fault diagnosis:The precision of 3D printing is influenced by many factors. One of the most important factors is the health of a 3D printer. Hence it is necessary to monitor the condition of components of 3D printer. SVM can be efficiently used for the fault diagnosis of 3D printers. A study [7] on fault diagnosis of delta 3D printer proposes Transfer support vector machine (TSVM) technique which is first of its kind.

A study [7] on fault diagnosis of delta 3D printer proposes Transfer support vector machine (TSVM) technique which is first of its kind. Figure 3 shows that in this hybrid approach, after preprocessing the attitude signals, data is divided into source domain and target domain. Next, cross-domain features are extracted from a source domain labeled-data and target domain unlabeled-data by performing transfer component analysis. Later SVM is used for classification using this new data as training data. This technique can better distinguish different fault conditions.Attitude se n sorStep .lChannel lChanne l 9Step 2"t- IF~•!!i' •"4 (_f • ~ ij III ►•~I-~ ! i : '1:)~flli ··II- 1i • ~~Ill Ill•• ~I!'~ ,:t -I .. 1 : :'8Dataset di.111.eus i o n a l 1·e du ctioorep JSource doma inT asgc t dorn a inS t ep 4T A n ...

Process maps:In addition to the sign of the decision function, its value plays an important role. [9] Has found that the value of decision function has a physical meaning which can be considered to generate a process map for the additive manufacturing process. This study uses SVM to classify the parts built by the EBM process in two classes ‘Flat’ or ‘Non-flat’. According to the study, the probability of obtaining an AM part with a flat surface increases as the value of the decision function increases. It is assumed that surface flatness reflects the energy balance between the input energy and energy loss and is related to internal defect generation. Thus, the value of decision function can be interpreted as a measure of porosity generation.

Thus, the value of decision function can be interpreted as a measure of porosity generation. This method uses a small amount of data to determine the process window and can be used to optimize the process parameters.218Random ForestsRandom forests are another supervised machine learning method used to build predictive models for both classification and regression problems. Random forest algorithm uses an ensemble of randomly generated decision trees to arrive at the best possible answer.Figure 5 depicts a typical decision tree. Decision tree creates a model to recursively subdivide the data. The aim while choosing every single division in the decision tree, starting from the root node, is to maximize the amount of information gain obtained by that division. Information gain is calculated by finding out the Gini impurity or the entropy gain for that division. Gini impurity is basically the probability (P) of incorrect classification of the ith class.

This algorithm works well with a completely different scale features, or a mix of the binary and continuous feature. However, the model tends to overfit and provides poor generalization performance [10]./~Fih@i -~Fiii@iIMU&IIMU&IFigure 5: Decision TreeRandom forests address this problem by creating an ensemble of decision trees. As shown in Figure 6, every random forest is built using two methods- bootstrap aggregating (bagging) and random subspace method. In bagging, bootstrap datasets are created which have the same size (n) as the original data set. The no. of these bootstrap datasets is the same as the no. of trees (B) in random forest. These datasets are created from random resampling of data with-replacement due to which datasets can have duplicate entries as well as missing entries. This process is called bagging. Next, to create a decision tree/ classifier, ‘m’ sub features are randomly selected out of ‘M’ possible features in the bootstrapped dataset.

This process is called bagging. Next, to create a decision tree/ classifier, ‘m’ sub features are randomly selected out of ‘M’ possible features in the bootstrapped dataset. In most cases, m=(cid:958)(cid:16). This is called a random subspace method.219(5)Input Training data: (X, Y)Select no. of features(M)Select no. of trees in RF (B)For,b= l to BCollect bootstrap data sample from training data• • • •Select specific no. of variables from bootstrap dataset(m)Run the model on all the points in that datasetrecord the output of decision treeOutput from Random Forest (Majority or arithmatic mean)CrossvalidationFigure 6: Flowchart of the Random Forest AlgorithmFor a random sample (Xb, Yb), for b= 1 to B, bagging and feature selection are carried out B times and a classification or regression tree is trained to have B possible outcomes.

The output of the random forest is either the majority of the outputs from all decision trees for classification problems or the arithmetic mean of the outputs of all decision trees for regression problems.Accuracy of random forests can be controlled by choosing the right number of features that are selected and correct depth for decision trees. Since each decision tree works on a different dataset due to bootstrap sampling, this ensemble method decreases the chances of overfitting [11]Applications of Random Forests in Additive ManufacturingBeing one of the most accurate learning algorithms, random forests algorithm has applications in defect detection, defect prediction which improves the part quality as well as helps in detecting the cyber physical attacks in additive manufacturing.1. Surface Roughness Prediction In [12] the algorithm is used to train the model that predicts surface roughness of the parts produced by the FDM process.

Surface Roughness Prediction In [12] the algorithm is used to train the model that predicts surface roughness of the parts produced by the FDM process. The study applies feature level fusion process for feature extraction from multiple sensor data such as the temperature of the table and extruder, vibrations of table and extruder, etc. to combine them into a single feature vector. This vector is input to the model which by using regression trees predicts the value of surface roughness. Also, models with individual sensor inputs are also generated. However, sensor fusion proves to be a more accurate method with the lowest cross validation error of 5.91%.2. Detection of attacks on cyber-manufacturing systems [13] attempts to use a vision-based system to detect intentional attacks on additive manufacturing processes, employing random forests algorithm.

Detection of attacks on cyber-manufacturing systems [13] attempts to use a vision-based system to detect intentional attacks on additive manufacturing processes, employing random forests algorithm. Due to malicious attacks on AM220systems, final products can result in defective infills without affecting the exterior and consequently can produce malicious defective parts without any warning. Random forest is used to accurately classify the in-process images parts as defective and non-defective in order to detect the anomalies during the process.Collection ol ca mera am! simulation imagesfor eadl image .. Extractior. ofplotting grayscal e valuesstatistical features.. Random forestalgorittim.. Output fromRa nclom Forest (Maj ority)..

Extractior. ofplotting grayscal e valuesstatistical features.. Random forestalgorittim.. Output fromRa nclom Forest (Maj ority).. CrossvalidationFigure 7: Procedure for image classification using the random forest algorithmFigure 7 shows the general procedure for image classification for detecting malicious attacks based on grayscale values of the images, mean, STD deviation, and no. of pixels with a grayscale value above a certain threshold are used as the input features for the random forest algorithm. For accurate feature extraction anomaly detection techniques are embedded with a random forest method which is proved to be advantageous as the detection accuracy is increased almost by 5% [14].Summary and ConclusionThe support vector machine algorithm works by generating a hyperplane that can divide the data into two or more classes. The hyperplane generated needs to be tuned in order to perfectly classify the data.

The hyperplane generated needs to be tuned in order to perfectly classify the data. This algorithm is successfully used for detecting part defects, diagnosing faults in 3D printers and generating process maps. The accuracy of this algorithm can be increased using different methods for converting input data from the input space to feature space such as transfer component analysis, combining data from multiple sensor modalities etc.Random forest algorithm proves to be another powerful classifier. It reduces the chances of overfitting and can be used in additive manufacturing as it can work with missing data and does not require scaling. It has been used in research on surface roughness prediction as well as in detecting malicious defects caused by the attacks on cyber manufacturing systems. Accuracy of this algorithm depends upon the no. of decision trees used and no.

Accuracy of this algorithm depends upon the no. of decision trees used and no. of levels of each decision tree.This review shows that, in additive manufacturing, machine learning basically has been applied for defect detection and prediction purposes. Table 3 summarizes the applications of SVM and RF in different additive manufacturing processes for different purposes.AimDefect DetectionSurface topologyPorosity PredictionProcessFDM, PBF FDM, EBM DLDMaterialABS, PLA, Stainless steel PLA, CoCr Alloy Ti-6Al-4VML technique SVMRF, SVM SVMReferences[6], [4] [12], [7] [5]221Fault Diagnosis in 3D printers Cyber Attack DetectionFDMFDMSVMRF[7], [8][13], [14]Table 3: Summary of applications of ML in AMFigure 8 shows the areas of the applications of both algorithms. While selecting an algorithm for any application there are different parameters that should to be taken into consideration.

While selecting an algorithm for any application there are different parameters that should to be taken into consideration. The input dataset, the method of data preprocessing as well as training the model, dataset taken as ground truth, computation power of the system, the time required for computation are some of the parameters. Thus, the choice of the algorithm is highly application specific which makes it difficult to compare between the two algorithms. For example, as shown in Figure 8, the only common area of research, under the scope of this review, which uses both SVM or RF is regarding the surface topology of the AM parts.

For example, as shown in Figure 8, the only common area of research, under the scope of this review, which uses both SVM or RF is regarding the surface topology of the AM parts. However, the performance of both algorithms cannot be compared as SVM is employed for classification and RF is used for regression purpose.Printers Fault DiagnosisSupport Vector MachineDefect Detection (Porosity, Inclusions)Process Map GenerationRandom ForestCyber Attack detectionFigure 8: Areas of application of ML algorithmsMore research is required in the metal AM processes as till now the majority of studies were conducted on the applications of SVM and RF in various fields of the FDM process. Within the scope of this review, it can be stated that SVM is preferred in comparison to RF. This review is limited to the study of only two supervised learning algorithms.

Within the scope of this review, it can be stated that SVM is preferred in comparison to RF. This review is limited to the study of only two supervised learning algorithms. Within the scope of the review, both algorithms prove to be highly accurate in the classification of parts as defective and non- defective. This makes them eligible for the application of detecting the generation of defects in real time.References:1. Gao, W., Zhang, Y., Ramanujan, D., Ramani, K., Chen, Y., Williams, C. B., … Zavattieri, P. D. (2015). The status, challenges, and future of additive manufacturing in engineering. Computer-Aided Design, 69, 65–89.doi:10.1016/j.cad.2015.04.0012. Gholami, Raoof, and Nikoo Fakhari. “Support Vector Machine: Principles, Parameters, and Applications.” Handbook of Neural Computation, 2017, pp. 515–535., doi:10.1016/b978-0-12-811318-9.00027-2.3. Kim, Felix H, and Shawn P Moylan.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperApplied Solvent-based slurry stereolithography process to fabricate high- performance ceramic earrings with exquisite detailsJia–Chang Wang1 Ph.D., Hitesh Dommati21 Department of Mechanical Engineering, National Taipei University of Technology; jcw@mail.ntut.edu.tw 2 College of Mechanical & Electrical Engineering, National Taipei University of Technology; iamhitesh89@gmail.comAbstract: This paper discusses the application of Slurry based Stereolithography additive manufacturing process in the fabrication of complex earring models without any support structures requirement, using High-Performance Ceramic (HPC) materials. The earring model chosen in this study is a Rose flower with blossomed petals. The petals have edge thickness in microns and extreme overhangs with a custom text and logo on the bottom.

The earring model chosen in this study is a Rose flower with blossomed petals. The petals have edge thickness in microns and extreme overhangs with a custom text and logo on the bottom. Using any other ceramic additive manufacturing process, it requires support structures to build this model. The support removal in such minute structures is not easy and not always successful. Using Solvent based Slurry Stereolithography (3S) process, models with the micron details and overhangs can be easily built. This is enabling the neat and clean post-processing procedure to maintain the exquisite details and also gain high surface quality. The ceramic material used in this application is alumina. With some additives, it will show in different colors like sapphire. The resultant flowers are vividly shown in white, pink, green, and blue.

The ceramic material used in this application is alumina. With some additives, it will show in different colors like sapphire. The resultant flowers are vividly shown in white, pink, green, and blue. In this study, it is also discussed about the slurry process, Stereolithography system, and proven applications of the 3S process.Keywords: Slurry based Stereolithography, High Performance Ceramics, non-self-supporting structures, earringsI.INTRODUCTIONAdditive manufacturing (AM) is a class of technologies that use Computer Aided Design (CAD) based layer-by-layer manufacturing process to construct the parts. Unlike conventional material removing process, AM works in complete contrary by adding layer upon layer. AM is considered as a driving force of “Third Industrial Revolution” since it has revolutionized the way we manufacture almost everything.

AM is considered as a driving force of “Third Industrial Revolution” since it has revolutionized the way we manufacture almost everything. There are multiple AM technologies that can shape High Performance Ceramics (HPC) to net-shaped parts by the resin consisting of structural materials and binders. Most common AM technologies that are used to for shaping ceramics are Material Extrusion (ME), Vat Photopolymerization (VP), and Powder Bed Fusion process (PBF) [1, 2, 3]. The ceramic AM process is classified into direct fabrication process and indirect fabrication process. In the direct fabrication process, the structure particles are fused when exposed to powerful laser beams and doesn’t require further thermal sintering process. The obtained surfaces of the parts through this process are rough because of the local thermal stresses due to temperature gradients. The Selective Laser Sintering (SLS), Selective Laser Melting (SLM) function as2790direction fabrication process.

The Selective Laser Sintering (SLS), Selective Laser Melting (SLM) function as2790direction fabrication process. Whereas, indirect fabrication techniques such as Stereolithography (SL), Fused Deposition Modeling (FDM) requires a three-step process, consisting of three dimensional printing, thermal debinding and Sintering. Having less internal stresses in this process, it is possible to fabricate complex geometries with sufficient structural strength and adequate mechanical properties. In photopolymerization process, the formation of the three dimensional parts happens on the layer-by-layer basis on a ceramic suspension in a photopolymer resin. The SL process uses a light source such as a projector which projects the light mask dynamically through a Digital Micromirror Device (DMD). The intensity of the light and exposure time are completely adjustable by the user depending upon the characteristics possessed by the resin.

The intensity of the light and exposure time are completely adjustable by the user depending upon the characteristics possessed by the resin. Repeating this step until the last layer, a green part is obtained. The green parts are cured ceramic parts but very soft in nature due to the presence of the binder within them. To remove the binder and strengthen the green parts, it is necessary to follow the thermal debinding and sintering processes. During debinding, the binder inside the green parts will evaporate and escapes through the surfaces incurring internal voids. While sintering, the structure particles bond with each other filling the void and causes shrinkage. The property of shrinkage of ceramics green parts during sintering is inevitable but, it can be controlled to a certain amount by altering the volume of binder. The occurrence of cracks in the green parts during the debinding and sintering process is a common phenomenon which is important to be avoided.

The occurrence of cracks in the green parts during the debinding and sintering process is a common phenomenon which is important to be avoided. It can be optimized by adjusting the rates of rising and fall of temperatures during debinding and sintering process.Conventional ceramic manufacturing methods can produce parts with high mechanical properties but they are not cost efficient for unit production or small scale production. They cannot produce high resolution objects due to the limitations of high tool wear during the grinding process [4]. Using AM technologies, it is possible to manufacture any high complexity geometrical structures in small scale at reduced costs and high resolutions. Yet, there is a limitation for topology optimization where it is difficult to fabricate ceramic objects which consist of non-self-supporting details.

Yet, there is a limitation for topology optimization where it is difficult to fabricate ceramic objects which consist of non-self-supporting details. For parts having details which are facing downwards to the surface called overhangs, they can be fabricated efficiently using AM techniques only if they persist above a critical value with respect to the build platform. Any part having details which do not meet the critical values, they are non-self-supporting parts and requires sacrificial support structures added to the overhangs in order to fabricate in an AM system. Langelaar.M proposed a Topology Optimization (TO) method that overcomes the limitations of the design process by implementing layer wise filtering and avoiding the unprintable objects from the design space resulting in fully self-supporting design [5]. It is not a common aspect for every part to be a self-supporting structure and falls within the critical value of overhangs.

It is not a common aspect for every part to be a self-supporting structure and falls within the critical value of overhangs. Parts with overhangs can be arranged in specific orientations to avoid sacrificial support structures following the TO techniques but, it is uncertain for every similar case. The critical angle of the overhangs that are typically found to be in the order of 450 [6]. Any part that has geometrical details which fall less than the critical angle, can be adjusted by changing the orientation or providing sacrificial support structures.

Any part that has geometrical details which fall less than the critical angle, can be adjusted by changing the orientation or providing sacrificial support structures. The addition of sacrificial support structures have clear disadvantages in terms of modifying the CAD file that may affect the mechanical properties, and consequently, it requires manual machining to remove them during post fabrication which increases the material and machining costs.No studies had supported fabrication of HPC’s using AM process without the requirement of sacrificial supports until Solvent based Slurry Stereolithography (3S) process was implemented by2791Wang J.C [7]. The 3S process can produce HPC parts by using a light source of a projector exposing photo masks on to the slurry layers from the top. This process attributes the novelty of conciliating any ceramic materials such as Alumina or Zirconia as structure materials with a recommended proportions of photopolymer agents and solvents.

This process attributes the novelty of conciliating any ceramic materials such as Alumina or Zirconia as structure materials with a recommended proportions of photopolymer agents and solvents. Also, the unmasked portion of slurry on the build surface acts as sacrificial support structures to the curing portion of the next layer where overhangs are present. Thus it doesn’t require any sacrificial support structures design for the overhangs. In this paper, the 3S process has been implemented and discussed in depth about its novelty of manufacturing complex geometrical structures which is an earring in this case.II. Materials and Methods2.1. Solvent based Slurry Stereolithography systemThe 3S system comprises of a DLP projector above a build platform making this a Top- Down system. The build platform moves downwards for every processed layer until the last layer.

The build platform moves downwards for every processed layer until the last layer. The layer coating is enabled by a scraping mechanism where a scarping blade distributes the deposited slurry evenly onto the platform and pushing out the unwanted portion of slurry off the platform. Once the solvent vaporizes, the DLP system exposes a masked light onto the surface of the slurry layer. The masked portion of the light over the slurry photopolymerizes and links to the previous layer.Figure 1: Solvent based Slurry SLA (3S) system2.2. Slurry with Ceramic suspensionThe slurry in this process is prepared by blending the main structural elements with a highly reactive photopolymer resin, methanol as a solvent and a dispersant. The structure element used in this study is Alumina (Al2O3). This forms a homogenous mixture by subjecting it to ball milling for one day using Alumina balls as grinding media.

The structure element used in this study is Alumina (Al2O3). This forms a homogenous mixture by subjecting it to ball milling for one day using Alumina balls as grinding media. The target viscosity of the slurry should be in the range of 2800Mpa - 3200Mpa or capable of spreading uniformly on the platform covering the build area evenly. An inorganic pigment of orange color is mixed in the slurry. It aids to absorb the blue light from the projector and suppress the light scattering over the slurry layers caused due to diffraction in order to control the overgrowth.27922.3.Manufacturing of non-self-supported structures HPC earrings using 3S process:To manufacture non-self-supported structures HPC earrings using 3S process, thefollowing steps are involved:1. STL file orientation and slicing 2. Green parts formation 3. Post processing 4. Post curing 5.

STL file orientation and slicing 2. Green parts formation 3. Post processing 4. Post curing 5. SinteringSTL file orientation and Slicing: The CAD file format commonly used in the AM technologies is STL which is a globally recognized format and recognized by every CAD software. The STL file of a Rose model (Opensource model, Thingiverse.com) considered in this study. The original model of the Rose is having a flat bottom as a base without any provision for the stud. It is as shown in figure 2(a). It is modified to by removing its base and applied a hole at the center. This hole is not a pass through but it can accommodate a stud. The modified rose model is as shown in figure 2(b). The structure also has a custom text ‘3DT Lab’ written on the bottom of a petal as shown in figure 2(c). This model has extreme overhangs with critical angle measured in negative values.

The structure also has a custom text ‘3DT Lab’ written on the bottom of a petal as shown in figure 2(c). This model has extreme overhangs with critical angle measured in negative values. The model is scaled to 13.44mm*13.36mm*10.10mm in length, width, and height such that the bottom hole can accommodate a stud. The edge thickness of the model’s petals are very minute and it is very difficult to fabricate by using any other AM system.(a)(b)(c) Figure 2: The Rose model: (a) Original Rose CAD model (b) Modified Rose CAD model (c) Customized text2793Green parts formation: A certain amount of slurry is deposited layer after layer onto the build platform as shown in figure 3(a). The deposited slurry is uniformly distributed onto the previous layer using a scraping tool. For every layer, the DLP system projects the sliced pattern as a mask on to the slurry surface for 1 second. The projected mask is based on a 405nm wavelength LED Full HD projector.

For every layer, the DLP system projects the sliced pattern as a mask on to the slurry surface for 1 second. The projected mask is based on a 405nm wavelength LED Full HD projector. The masked portion of the slurry gets solidified once exposed. The unmasked portion of the slurry remains soft and acts as a support structure for the proceeding layer. When the next layer is exposed to the corresponding lighting mask, that portion is linked with the previous layer and forms a three dimensional part. In this study, the objects are printed at the layer resolution of 20µm. Figure 3 depicts the DLP projecting a light mask to cure a layer.(a)(b)Figure 3: Green parts formation: (a) Slurry deposition (b) Exposing light mask for curingPost processing: Once the green parts are fully formed, it is necessary to take the build platform out of the system and clean the uncured slurry off the surface carefully with water. Figure 4 represents the post processing of the cured green parts.

Figure 4 represents the post processing of the cured green parts. In cases of models with complex geometries and soft portions, it is difficult to manually clean the slurry in the complex regions. Hence, the green parts are cleaned thoroughly using an ultrasonic cleaning equipment.Figure 4: Post processing using waterPost curing: The obtained green parts are the outcome of an incomplete reaction of photopolymerization in the 3S system. They are very soft and deformable in nature. To complete the reaction, they are subjected to 405nm light to post cure for 10 minutes. During post curing, the soft parts lose deforming nature and obtain firmness. Yet, the firm parts do not completely attain the properties of the HPC’s. They still carry the binder within them causing to be fragile. The following figure 6 depicts the outcomes after post curing.

Yet, the firm parts do not completely attain the properties of the HPC’s. They still carry the binder within them causing to be fragile. The following figure 6 depicts the outcomes after post curing. The dimensions are approximately similar to the values of the CAD model.2794Figure 6: Roses after post curingDebinding and Sintering: The debinding is the step which thermally heats and evaporates the binder present in the post cured green parts. The evaporated binder escapes through the nearest surface of the parts. An abrupt increase in temperature causes non-uniform conduction of heat resulting in cracks due to the non-uniform evaporation of the binder [8]. The rose model has thin features and which helps the binder to exit from the model without causing any cracks. Figure 7 depicts the Thermogravimetric analysis (TGA) test of the binder used in this process. Following the plot, the green parts are subjected to raise in temperature from 40- 3500C for 2 hours.

Figure 7 depicts the Thermogravimetric analysis (TGA) test of the binder used in this process. Following the plot, the green parts are subjected to raise in temperature from 40- 3500C for 2 hours. At 3500C, the binder starts to evaporate faster. Hence, the raised temperature is increased to 6000C in next 5 hours in which most of the binder is evaporated and leaves their spaces empty.Figure 7: TGA test of the binder used for thermal debindingThe debinding process is then followed by sintering process where the parts are heated from 6000C to 16000C. The atoms in the green parts diffuse across the boundaries of the particles, fusing the particles together creating one solid piece and fills the voids. The contraction of the particle causes the parts to shrink. The sintering cycle starts heating the parts at 6000C to 16000C in 3 hours of time and cools down from 16000C to 450C in 6 hours.

The contraction of the particle causes the parts to shrink. The sintering cycle starts heating the parts at 6000C to 16000C in 3 hours of time and cools down from 16000C to 450C in 6 hours. Figure 8 represents the roses obtained after sintering.2795Figure 8: Resultant models of the roses after sintering2.4.3D printing in different colors using 3S process:The 3S process enables to 3D print the HPC’s in various solid colors like blue, green, pink, etc. By adding coloring agents like Nickel Oxide, Cobaltous oxide, Chromium oxide, while preparing the slurry, green, blue and pink colors are obtainable after sintering. Figure 9 depicts the resultant models of the rose which are produced after sintering.Figure 9: 3D printed Roses with Alumina using 3S process in different solid colors2.5.Making of Rose earrings:The bottom hole on the resultant models is used to fix a stud which fits within. Once the stud is fixed in, it is glued to permanently fix.

Once the stud is fixed in, it is glued to permanently fix. Figure 10 depicts the roses having fixed with the studs to complete as an earring.2796Figure 10: A stud fixed to the bottom hole of the roseIII. Results and DiscussionThe resultant parts are now dense ceramic rose earrings which possess complete properties of any ceramic material. These objects are hard but brittle in nature. The overhangs of the structure are strong enough to have no cracks at any portion of the part. The edges are well constructed and maintained firmly even after sintering. The resolution of the parts are analyzed using 3D Macroscope (Keyence VR – 3000 series) and are described in the following. Shrinkage: The HPC’s are unavoidable of shrinkage when sintered at high temperatures. By sintering, they tend to shrink uniformly and become strong and dense components [9]. The 3S process implemented by Wang J.C. shows 20% shrink rate in alumina built parts.

By sintering, they tend to shrink uniformly and become strong and dense components [9]. The 3S process implemented by Wang J.C. shows 20% shrink rate in alumina built parts. To measure the shrinkage, a standard rectangular model with round shaped edges and two holes on both ends is printed using 3S process and its dimensions are measured by a Macroscope equipment. Three sample models are 3D printed and one of the parts which is post cured is as shown in the figure 12(a). The figure 12 (b) depicts the 3 sintered samples.

Three sample models are 3D printed and one of the parts which is post cured is as shown in the figure 12(a). The figure 12 (b) depicts the 3 sintered samples. The before sintered and after sintered samples dimensions are evaluated by the Macroscope equipment and the details are mentioned in the Table 1.(a)(b)Figure 12: (a) Sample before sintering (b) Samples after sintering2797SamplesSample 1Sample 2Sample 3Pre- sintered27.827.8727.79Sintered232322.94Linear Shrinkage20.87%21.17%21.14%ResultsAverage 21.06% sample standard deviation 0.165 %Table 1: Length dimensions of 3 samples before and after sinteringThe length of the sintered samples of 1,2 and 3 obtained are of 27.8mm, 27.87mm and 27.79mm. The length is measured by considering the midpoint of the part that can be obtained by drawing a line from the center points of both the holes. The same samples after sintered have the length of 23mm, 23mm and 22.94mm.

The same samples after sintered have the length of 23mm, 23mm and 22.94mm. The differences in the lengths gives the linear shrinkage values for which the sample 1 has the shrinkage value of 20.87%, sample 2 has 21.17% and sample 3 has 21.14%. The average linear shrinkage value is 21.06% and the sample standard deviation is of 0.165%.Resolution: The resolution of the parts can be divided into two categories which are pixel resolution and layer resolution. The pixel resolution gives the quality of the parts in X and Y axis. This is completely dependent on the projector used. The projector used is a Full HD having 1920*1080 pixels in X and Y directions and the setup of the dots size per pixel is 40µm/pixel. This gives the working area size in total length and width of 76.8mm and 43.2mm. Secondly, layer resolution exhibits the parts quality in the Z axis.

This gives the working area size in total length and width of 76.8mm and 43.2mm. Secondly, layer resolution exhibits the parts quality in the Z axis. The layer resolution is dependent on the structure particle size and the minimum displacement and repeatability the Z axis motor can achieve [10]. The d50 of the grain size of alumina particles used in this study is 0.39µm and the Z-axis displacement is 20µm. If calibrated the projector carefully 20µm resolution of the parts can be obtained.Discussion:By observing the cross section thickness of the above segment shown in figure 12(a), it is explainable about the 3S process and its system combined are able to build fine features without requiring sacrificial supports. The hardness of the profiles is dependent on the ceramic material which is alumina in this case. The hardness measurement studies of the alumina components fabricated by 3S process conducted by Wang.

The hardness of the profiles is dependent on the ceramic material which is alumina in this case. The hardness measurement studies of the alumina components fabricated by 3S process conducted by Wang. J.C indicate they are 98% dense and have a tensile strength of 327 MPa, flexural strength of 472 MPa [7]. The strength values vary according to the structure component added in the slurry. The build volume of this system is dependent on the projector’s resolution and the pixel size. The build volume can be increased by increasing the pixel size from 40µm to 50µm but, the quality of the parts degrades proportionally. Similarly, by decreasing the pixel size from 40µm to 30µm, the quality of the build parts aggrades but, the build size of the system is decreased proportionally. The shrinkage after sintering accounts in resulting2798the finest resolution of the parts of 32µm.

The shrinkage after sintering accounts in resulting2798the finest resolution of the parts of 32µm. Due to the thin features in the Rose model and temperature control for the debinding and sintering, no cracks were observed.During the curing process, one layer of the binder molecules is linked with the previous layer due to the light energy. This energy is directly proportional to the time. The light exposure time is adjusted to its optimum value by performing time to energy experiment where parts are built at different time values starting from 0.1s to 2.0s. The parts exposed for 0.1s to 0.8s resulted in poor curing and very soft which were broken off while post processing. The parts exposed for 0.9s to 1.1s provided good results whereas the parts exposed for greater than 1.1s were shown overgrowth due to over-curing and resulted in protruded structures causing the scraping blade to scrape them off during printing itself.

Hence, the optimum and mean exposure value of 1 second is considered to fabricate alumina parts.IV. CONCLUSION This work presents the fabrication of non-self-supporting high complex geometrical structures without any requirement of sacrificial support structures using Solvent based Slurry Stereolithography process. A rose model with critical overhangs is considered as the CAD model and modified it to an earring. Using alumina as the structure material, the slurry is prepared with appropriate proportions of photopolymer as a binding agent, dispersant, and methanol as solvent. The 3S system photopolymerizes every layer with lighting mask to cure the layers and forms fully cured ceramic green parts which are later subjected to post processing, post curing and sintering according to the standard procedure. The green parts have pertinent features of dimensions and accuracy of the CAD file. A shrinkage rate of 21.06% with a standard deviation of 0.165% is observed in the post sintering.

The green parts have pertinent features of dimensions and accuracy of the CAD file. A shrinkage rate of 21.06% with a standard deviation of 0.165% is observed in the post sintering. The resultant sintered parts are fixed with studs to enable to use as earrings. The details and features of the models are steady, strong and aesthetically pleasing. It is conclusive that the 3S process enables to build any non-self-supporting structures without any sacrificial structures at high resolution with accurate dimensions. The shrinking volume has to be considered as a parameter that has to be calculated to scale up the CAD file before printing, such that the final parts meets the desired dimension.ACKNOWLEDGEMENTSThe authors of this paper would like to thank the support by the Ministry of Science and Technology, Taiwan, R.O.C. under Grant no. MOST 105-2221-E-027-036-MY2.REFERENCES1. Scheithauer, U., Schwarzer, E., Richter, H.J. and Moritz, T., 2015.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperApplied Viscous Thread Instability for Manufacturing3D Printed FoamsB. Emery1, D. L. Revier2,3, J. I. Lipton2,31. Department of Physics, University of Washington, Seattle, Washington 98195, USA 2. Paul G. Allen School of Computer Science & Engineering, University of Washington, Seattle, Washington98195, USA3. Department of Mechanical Engineering, University of Washington, Seattle, 98195, USAAbstractTraditional foams are fabricated via stochastic chemical processes that yield homogeneous material properties. Foams can exhibit a wide range of material properties by varying process controls allowing them to be used in many industrial and commercial applications. Previously, additive manufacturing could only produce foam approximations in the form of traditional lattice infill.

Previously, additive manufacturing could only produce foam approximations in the form of traditional lattice infill. Our work employs viscous thread printing (VTP) of thermoplastic polyurethane (TPU) on a fused filament fabrication (FFF) printer, exploiting the semi-viscous nature of extruded filament to coil producing a new type of printed foam. Specimens were tested under compression to determine uniformity along principal axes and behavior under strain when compared to infill patterns, such as grid and cubic.

Specimens were tested under compression to determine uniformity along principal axes and behavior under strain when compared to infill patterns, such as grid and cubic. This work establishes that VTP as applied to TPU can be used to manufacture programmable stiffness foams as a function of density, suited to a variety of needs and should be considered as an alternative to traditional foams and other printed lattice geometries.Keywords: Foams, Viscous Thread Instability, Viscous Thread Printing, CellularStructuresIntroductionAdditive manufacturing (AM) processes, also called “3D printing”, are commonly used to fabricate porous cellular structures for a variety of applications from bio-printing for cellular scaffolding and medical implants [1,2] to mechanical applications for weight reduction as well as optimizing static and dynamic mechanical properties [3].

Traditionally, printed cellular structures are fabricated in one of two ways: 1) explicitly defining the geometry though design of the cells themselves or 2) implicitly through tool path processing such as volume-filling paths commonly called “infills.”The explicit definition of cellular structures requires printer resolution to be much finer than that of the cellular structure unit cell, usually by at least one order of magnitude or more [3– 6]. Furthermore, these geometries require design of complex cellular structures though defined cellular structures such as octet truss unit cells or through topology optimization [7–9], in which the design space is tightly bounded or arbitrarily large respectively. Ultimately, cell size and design are significant limitations to an explicitly defined cellular geometry for printed objects.Conversely, implicit cellular structure definition (i.e.

Ultimately, cell size and design are significant limitations to an explicitly defined cellular geometry for printed objects.Conversely, implicit cellular structure definition (i.e. an infill algorithm) is not sufficient for engineered materials because they often rely on “black box” algorithms, are geometric in33nature, and tend to be anisotropic [10]. Infills with back-and-forth toolpath patterns such as “grid” or more sophisticated geometric definitions such as “cubic” can only be reasonably controlled through one variable, the infill percentage, and furthermore must remain consistent or rely on discrete “steps” to incrementally increase density throughout the geometry, disallowing any possibility of material property gradients.

Ultimately, implicit cellular definition algorithms are adequate in many simple use cases, but are insufficient from an engineered materials perspective.In this work we extend a printing technique known as Viscous Thread Printing (VTP) [11] for use with thermoplastic polyurethane (TPU) on a common Fused Filament Fabrication (FFF) printers. We generate cellular foam structures at resolutions that are comparable to the native resolution of the printer by utilizing the phenomenon of Viscous Thread Instability (VTI). VTI is a well characterized phenomenon in 1D and 2D [12–15] that is familiar to anyone who has seen honey cascade from a spoon onto toast. The bending and coiling pattern that the viscous fluid makes as the thread traverses the surface is simple, yet well-controlled phenomenon in 1D.

The bending and coiling pattern that the viscous fluid makes as the thread traverses the surface is simple, yet well-controlled phenomenon in 1D. Extending to 2D and 3D, VTI becomes much less predictable or controllable; however, can still result in a wide variety of coiling patterns classified into shapes such as W’s, meanders, alternating loops and translating coils [13,15] which can be exploited to generate 3D structures when stacked.In this paper we demonstrate that VTP of TPU results in an implicitly defined, open-celled, foam structure with uniform stiffness along principal axes. We focus on translating and alternating coils since they produce the greatest number of interconnections between strands. We show that by manipulating simple print parameters we can print single-layer, non-woven textiles on a standard FFF 3D printer.

We show that by manipulating simple print parameters we can print single-layer, non-woven textiles on a standard FFF 3D printer. When stacked layer upon layer the results are 3D open-celled foams with programmable stiffness as a function of density that can vary over multiple orders of magnitude.TPU VTP Process CharacterizationTPU behaves as a viscous liquid in its semi-molten state when printed using an extrusion- based process such as FFF. The key parameters to determine the coiling behavior of the TPU are the dimensionless velocity variable V* and the dimensionless height variable H*,𝑉∗ =𝐻 ∗ =𝐹 𝐶 𝐻 α𝐷where F is the linear translation speed of the printhead, C is the linear velocity of the material being extruded, H is the print height, D is the nozzle diameter and α is die swell, a material and process dependent constant. Each of these variables can be controlled either explicitly or implicitly directly through g-code commands.

Each of these variables can be controlled either explicitly or implicitly directly through g-code commands. Together, the V*,H* combination have a dominant effect in determining the type of coiling behavior exhibited as well as the spatial frequency of the coiling respectively [12,13,15].There are two main regimes of coiling that are of interest for VTP foams: translating coiling and alternating coiling. Translating coiling results from a specific combination of V*,H* to produce a continuous path of overlapping loops of material. Translating coiling is most interesting because it allows for the greatest number of interconnections of a single thread producing a structure that is porous and homogenous in macro structure. Alternating coiling may be used for34VTP foams but is less desired because it produces fewer interconnections and is, in general, a less regular structure. Finally, one other significant regime is accumulation (i.e. over-extrusion).

Finally, one other significant regime is accumulation (i.e. over-extrusion). Accumulation is not desired because the thin nature of the thread is lost and instead produces solid structures more similar to conventional FFF parts than foams [15].1D/2D CharacterizationFirst, we established preliminary bounds on V*/H* for TPU by printing spreads of 1D lines to establish the range of coiling behaviors and frequencies able to be produced as shown in Figure 1. Initial tests were performed using already established V*/H* bounds for prior work on similar materials [11] and then were expanded to test full theoretical bounds for varying coiling instability domains. We determined for TPU that V* within the range of 0.12-1.0 and H* >5 for 1D lines produced the variety of coiling behaviors and densities that would allow for control of foam density.

We determined for TPU that V* within the range of 0.12-1.0 and H* >5 for 1D lines produced the variety of coiling behaviors and densities that would allow for control of foam density. For V* < 0.12 we observed over-extrusion, also called accumulation, that would have required unrealistic amounts of filament to be extruded, resulting in a lack of any significant coiling. For V* > 1 no coiling behavior was observed and the TPU had difficulty sticking to the build plate in many cases. For H* < 5 we observed a very high coiling frequency bordering onFigure 1 - V*/H* sweep from 0.1-1 and 5-10 respectivelyaccumulation and for H* > 10 the TPU had difficultly sticking to the build plate for many V* values greater than 0.45 making regular coiling difficult to control.35For characterization in 2D, coil meshes are formed by continuously extruding material in a zig-zagging toolpath pattern as seen in Figure 2.

The length of any side of the zigzag is defined as L, while the spacing between continuous L moves is λ. In this case L defines the overall geometry but λ contributes directly to the density of the mesh in addition to the consistency of coiling. λ must be empirically determined so that the coiling behavior can be properly controlled to ensure good overlap between passes of coils while also programming the density of the mesh. In all, the translation from 1D to 2D coiling behavior is straightforward with the major exception that as coils begin to overlap the resulting new coil becomes unstable and unpredictable. For any isolated 1D line the coiling behavior will reach a metastable state that can easily be perturbed by a non-uniform build surface or a change in velocity.

For any isolated 1D line the coiling behavior will reach a metastable state that can easily be perturbed by a non-uniform build surface or a change in velocity. Once perturbed, the coiling behavior will be temporarily destabilized until the printed line is no longer being influenced and is allowed to reestablish its metastable state.3D CharacterizationFigure 2 - Zigzag toolpath pattern to produce a 2D meshIn order to achieve a well-formed cube, the parameters V*, H* and λ are tuned to program the density for a one-layer mesh. The single layer thickness Δz is then measured by hand to determine the necessary z-axis increment to build successive mesh layers on top of each other to achieve the desired density 3D foam cube. This is affected by many factors such as how much parallel coils overlap, how much the loops of a coil overlaps with the next, and how flat the build surface is.

This is affected by many factors such as how much parallel coils overlap, how much the loops of a coil overlaps with the next, and how flat the build surface is. A constant Δz is used to achieve an approximately constant H* value at a given layer to maintain as consistent of a coiling pattern as possible. The zig zag pattern is then rotated for each layer by 90o resulting in a crosshatch toolpath to generate the coiled layer.Foam homogeneity is very sensitive to the measurement of Δz. Significant non-uniformity in foam density can occur as the foam is built up even for a small error in Δz. For overestimates36of Δz, the density of the foam will decrease in the vertical direction. This is due to the increasing effective H* as the nozzle height above the foam is increasing resulting in lower frequency coiling, fewer coil overlaps, and lower layer density.

This is due to the increasing effective H* as the nozzle height above the foam is increasing resulting in lower frequency coiling, fewer coil overlaps, and lower layer density. Underestimating Δz will result in accumulation as the distance between nozzle and foam will continually decrease, shrinking H*, and increasing the density of the foam in the vertical direction. A properly calibrated Δz will result in a part that is nearly homogeneous, properly proportioned, and the correct height.There are two methods of varying foam density: varying the V*,H* combination or sweeping λ. Controlling V* and H* allow for precise control of coil behavior and frequency; however, manipulation of these variables requires recalibration of λ and Δz to achieve a well- formed cube and specific density. This means that while a range of acceptable densities is achievable through direct V* and H* manipulation, at this time it is very difficult to specify a density without many trial and error experiments.

The second method of programming foam density is by varying λ. As λ is swept it goes from producing significant overlap of coils to significant “dead space” between coils for a constant V*,H* combination. There is an inverse relationship between density and λ and so density can be targeted with better precision; however, for very small λ there is so much accumulation of coils on top of each other that a layer can fused into a single continuous printed piece without coiling behavior. For a very large λ relative to coil diameter the coils operate more as a loose scaffolding pattern than an open-cell foam structure.As a rule, the density of foam is directly proportional to the quality of fine details and sharp edges. Foams with relative densities 20% or lower tend to lose their shape and have rounded features, where foams with relative densities >50% begin to have significant accumulation issues. Figure 3 shows four small cubes printed at increasing density from left to right.

Figure 3 shows four small cubes printed at increasing density from left to right. The foam on the left was produced with very large coils to achieve a low density and so lacks the fine corners seen in other cubes. Conversely, the foam on the right was printed with such a small λ that accumulation caused high intra-layer adhesion and extremely poor inter-layer adhesion so that successive layers did not stick to each other well and result in an “onion” effect where layers peel apart easily. While the feature resolution of this geometry is very good, it clearly does not operate as a foam.Figure 3 - Increasing density foams approximately 15%, 25%, 40%, and 60% left to right respectively37Results and DiscussionSpecimen Fabrication & TestThe filament used for all trials was NinjaTek NinjaFlex TPU (85A) with a density of 1.19 g/cc and a measured a die-swell (α) of 1.18.

All specimens were printed at a nozzle temperature of 235°C with a build plate temperature of 40°C, and a feed rate (F) of 60 mm/min. Each specimen was printed in a cubic formfactor with nominal measurements of 50x50x50 mm3. For every unique set of parameters, two specimens were printed to test the both the vertical and horizontal mechanical properties of each specimen. VTI foam specimens were prepared by varying V*/H* and/or λ at nominal densities of 15%, 25%, and 40% as outlined in Table 1. Cubic and grid infill specimens were printed for comparison at the same nominal densities of 15%, 25%, and 40%. These two traditional infill patterns were chosen because they are both highly prevalent in all forms of additive manufacturing and each demonstrate distinct cellular properties for our foams. Grid infill was anticipated to be highly non-uniform in stiffness along principal axes and cubic infill more uniform.

Grid infill was anticipated to be highly non-uniform in stiffness along principal axes and cubic infill more uniform. All objects were printed without top or bottom layers nor walls, leaving only the inner geometry to bear all load placed upon the specimen.Figure 4 – TPU foams. From left-right, top-bottom: Samples 11, 18, 19, 24 and 9.3Figure 5 - Cubic infills (a) and grid infills (b) with nominal densities 15%, 30%, 40% from left to right in each photograph38Table 1 - Foam fabrication variables and resulting densities normalized to TPU intrinsic density 1.19 g/cc.Sample No. 24 9.3 11 18 19Nominal Density Ratio15% 15% 25% 25% 40%Actual Density Ratio 12% 18% 29% 24% 36%V*0.23 0.12 0.23 0.23 0.23H*7.5 15 7.5 7.5 7.5Δz (mm)0.9 2 1.3 1.1 1.2λ (mm)6.5 2.5 1.65 2.5 1.35All specimens were tested under compression vertically and horizontally as compared to original print direction.

The force was applied by a 68SC-2 Instron machine, with test cell load limit of 1750N for 5 consecutive trials per specimen. Measurements were averaged together and a linear fit was determined for small strain applied in conjunction with stress as calculated by cross sectional area to determine the Young’s modulus of each specimen. Density ratio was calculated as a ratio of measured density and the given manufacturer’s density, where measured density was determined via measured mass over the product of measured length, width, and height of each specimen.Figure 6 – (a) Vertically oriented 15% dense grid infill buckling under compression and (b) specimen number 18 under compressionResultsFigure 7 shows the Young’s modulus of the printed foams compared to those of the printed grid and cubic infill patterns.

All VTP foams can be well fit onto an exponential trendline with significantly lower Young’s moduli on average than the infill specimens, with the notable exception of the horizontal grids. Tested specimens also follow a similar trend of horizontally oriented specimens generally demonstrating lower modulus values than vertically oriented specimens, likely due to the inherent layer orientation resulting from the FFF process.The grid infill specimens were found to be highly non-uniform in stiffness along principal axes. This is due to the internal structure of grid which may be oriented to exhibit greater value of Young’s modulus when vertically orientated as opposed to the significantly lower modulus of the horizontal orientation. When vertical the open cells of the grid form a very stiff structure that exhibit a greater Young’s modulus than any other geometry of similar density; this trend continued for all other densities tested.

This structure would behave in two distinct phases of deformation, prior to buckling and after buckling. Prior to buckling, the deformation resulting from load would39be distributed evenly over the entire structure. The original structure will be mostly maintained until a critical force is reached, at which point the structure will buckle and fold inward as seen in Figure 6(a). From this point on, the specimen’s stiffness would significantly decrease until densification. When horizontally oriented, grid exhibited values of Young’s modulus that were significantly lower than any other infill pattern and was similar to values of the foams. Therefore, it can be concluded that the stiffness of grid infill is determined by density as well as orientation.Foam & Infill Stiffness) a P k (100001000s ulu d o M100s ' g n u o Y101 10%15%20%25%30%35%40%45%Horizontal VerticalFoamsGridDensity RatioCubicExpon.

(Foams)y = 1.3873e17.435xFigure 7 - Foam and infill Young's modulus values. Grid infills tend to be highly non-uniform in stiffness along principal axes between vertical and horizontal testing, while cubic infills tend to be more uniform in stiffness along principal axes. Foam patterns also tend to be be uniform in stiffness along principal axes but with significantly lower Young’s modulus as compared to the infill patterns. An exponential trendline can be drawn for the foam’s with R2 = 0.9846. Full values listed in Table 2Cubic infill specimens were found to be highly uniform in stiffness along principal axes. This is due to the internal structure of the cubic infill being symmetrical along orthogonal axes. As such the stiffness of cubic infill specimens when vertically oriented were comparable to the stiffness of an equally dense cubic pattern when horizontally oriented.

As such the stiffness of cubic infill specimens when vertically oriented were comparable to the stiffness of an equally dense cubic pattern when horizontally oriented. Regardless of orientation the values of Young’s modulus for the cubic infills were consistently less than the values of the vertically oriented grid while also greater than the values of horizontally oriented grid and foams. While it is possible that air trapped inside the closed cellular structure could contribute to artificially high modulus values (due to compression of the air), we believe this is a negligible source due to the unlikelihood that layer adhesion of TPU is sufficiently airtight.

Thus, it may be concluded that the only characteristic relevant to the stiffness of cubic infill is density.Foams manufactured by sweeping λ as well as foams created by varying V*,H* were found to be predominantly uniform in stiffness along principal axes, where V*,H* varied foams exhibit less difference in Young’s modulus than λ modulated foams. Despite this finding, generally the foams were equally uniform or better than any other geometry tested. During testing it was observed that when horizontally oriented the face composed of the first layer would deform differently than the other exposed faces. This is due to the fact that the first layer is the most well- formed of all the layers of the foam since its build area was the most flat and consistent. This results in a greater stiffness than other layers, and therefore this layer will buckle where other faces will expand uniformly outwards. As seen in Figure 7 all tested foams can be well fit onto an40exponential trendline.

As seen in Figure 7 all tested foams can be well fit onto an40exponential trendline. The goodness of this fit can be best approximated by the R2 value of 0.98. Therefore, we believe this fit very closely approximates the trend at which all foams printed in this manner will follow, allowing us to use this trend to create a framework for converting density into input variables (V*, H*, and λ). From this, it can be inferred that the most significant characteristic of a printed foam’s stiffness is its density.ConclusionsUsing viscous thread printing (VTP) of thermoplastic urethane (TPU) we successfully manufactured parts that approximate the stochastic internal structure of an open-celled foam with programable stiffness using standard fused filament fabrication (FFF) equipment. These foams are customized using dimensionless variables V* and H* to define coiling behavior as well as measured variables λ and Δz to calibrate a homogeneous internal geometry.

These foams are customized using dimensionless variables V* and H* to define coiling behavior as well as measured variables λ and Δz to calibrate a homogeneous internal geometry. By modulating V* and H* or by modulating λ, the densities of these foams may be tailored to many potential needs or scenarios; however, when calibrating λ or Δz both variables will affect the overall density of the resulting part. Therefore, a balance must be reached between accuracy of detail and accuracy of density. The stiffness of each specimen was determined according to deformation under load. While under such load, our foams were shown to be highly uniform in stiffness along principal axes and followed an exponential Young’s modulus trend with respect to density, allowing engineering specific stiffness through simple control of the input parameters.

By comparing our foams to conventional infill patterns, we have shown that our foams possess pic properties similar to cubic infill along principal axes with stiffness most similar to horizontally oriented grid infill. As such our foams fill a unique niche leveraging the properties of some infill patterns while balancing their disadvantages where no standard infill can achieve both simultaneously.AcknowledgementsThis work was made possible by a gift from the Ford Motor Company as well as by a grant from the NASA Space Grant Initiative and scholarship support from the Mary Gates Endowment. We thank the Murdock Charitable Trust, for their support through grant 201913596.41References1. Derby B. Printing and prototyping of tissues and scaffolds. Science. 2012;338: 921–926.2. Melchels FPW, Domingos MAN, Klein TJ, Malda J, Bartolo PJ, Hutmacher DW. Additive manufacturing of tissues and organs. Prog Polym Sci. 2012;37: 1079–1104.3. Chu C, Graf G, Rosen DW.

e Approach of Complex Insert Packaging Fabrication in Stereolithography Y. Y. Chiu* and J. H. Chang++ Department of Industrial and Commercial Design, Oriental Institute of Technology,58, Sec. 2, Syh-Chuan Rd., Baan-Chyan, 220, Taipei, Taiwan, R.O.C.E-mail: petecyy@ms19.hinet.netAssociate Professor, Department of Industrial and Commercial Design, Oriental Institute of Technology, Baan-Chyan, Taipei, Taiwan, 220, R.O.C.++Research Assistant, Department of Industrial and Commercial Design, Oriental Institute of Technology, Baan-Chyan, Taipei, Taiwan, 220, R.O.C. Reviewed, accepted August 4, 2004AbstractThe approach of complex insert packaging fabrication in stereolithography is studied in thepaper.

Reviewed, accepted August 4, 2004AbstractThe approach of complex insert packaging fabrication in stereolithography is studied in thepaper. There are many difficulties being overcome, such as the dispensing shadowing problemcaused by the geometric shape of insert, the polymer feeding problem caused by greater insertheight than the layer thickness, and the positioning problem of insert that leads to instability ofpackaging. These drawbacks led to unsatisfactory results of the insert packaging instereolithography. In order to solve the problems, a new method of complex insertspackaging fabrication in stereolithography is proposed in this paper. Based on the geometricinformation, function and assembly direction of the inserts, the packaging approach is developed.The approach proposed in this paper has been verified by experiments. It brings considerablecontributions to the application of insert packaging in stereolithography.

It brings considerablecontributions to the application of insert packaging in stereolithography. It is also favorable tothe improvement of insert packaging efficiency and assembly fabrication.Keywords: Rapid Prototyping, Solid Freeform Fabrication, Insert Packaging, Insert1. Objectives and Motives: Solid freeform fabrication (SFF) or rapid prototyping (RP) has displayed amazing potentials in the realms of product design and rapid manufacturing. From the literatures of the recent years, it is known that the future research trend of RP is small-sizing of rapid prototyping objects [1] and the manufacturing of assembling objects having complex functions.

This paperaims at solving the problems and proposing a new processing method of insert packagingencountered in SLA’s manufacturing of complex shape, functional mechanical structure andassembling objects.In the early stage of RP development, it tends to be applied in the construction of original mold in the area of industrial design. However, in recent years, it tends to develop towards themanufacturing of functional assembling objects.Under an acceptable tolerance, themanufactured parts carry good precision, excellent surface precision and high flexibility. This572enables a total process of manufacturing the finished product, without going through additionalassembly [2] [3]. Furthermore, it can also establish the procedure that the parts of differentmaterials are all embedded in the simple-structured RP object [4], such as embedding the parts ofsensor, controller and bearing etc. into RP before the manufacturing of RP or during itsfabrication.

into RP before the manufacturing of RP or during itsfabrication. Kataria [5] et al. explored the method of using photo polymer (SLA) to construct complexmechanism. Focusing on the various processing features of SLA, such as the design of fixture,the interference problems of recoating and insert, geometric shape of insert, positioning andfixture mechanism, and geometry tolerance etc., they proposed a series of process planningmethods, and developed a multiple-shaft SLA machine [6], exploring into the RP processingrealm in times of manufacturing the complex assembling object. However, the problems ofdispensing shadowing, recoater interference, positioning and fixture etc. still limited the workability for SLA to manufacture assembling object and complex inserts. In order to avoid obstructing or damaging the functions of flexible part, Chad A. Moore et al.

still limited the workability for SLA to manufacture assembling object and complex inserts. In order to avoid obstructing or damaging the functions of flexible part, Chad A. Moore et al. [7] studied the method to embed flexible part into multi-material prototype, maintaining theshape of the flexible part in the process, and accurately controlling the geometric shapes in the contacting area of different materials. Finally, they proposed a series of processes.According to the theory that SLA process has to undergo post-cure processing treatmentafter the product is completed, this paper proposes “online insert embedding process.” Byusing the pre-analyzed and pre–designed component, support and cavity, the insert is positioned,and then photo polymer is poured in. Afterwards, post curing processing treatment is made tosolidify the required part in order to meet the objective of embedding in specific work-piece.2.

Afterwards, post curing processing treatment is made tosolidify the required part in order to meet the objective of embedding in specific work-piece.2. Approach The online insert embedding process proposed by this paper is to solve the problems currently encountered by SLA when manufacturing the assembling object, complex geometricstructure and functional structure. Therefore, algorithm of support of structure with geometric features is proposed, and furthermore, “online insert embedding process” is developed. It ishoped that the limitation of embedding insert in geometric shape in the SLA process can beeliminated. Focusing on the procedures of online insert embedding process and the theory ofalgorithm of support of geometric feature structure, detailed introduction and investigation areelaborated herein.Figure 1 shows the flowchart of SLA process. Among the various RP machines, SLAprocess is the one with the highest precision.

Among the various RP machines, SLAprocess is the one with the highest precision. However, due to the problems of laser system,recoating system and the need to construct supporting material, lots of obstruction have beenencountered during online embedding of work-piece.The procedures of online insert embedding process proposed by this paper are as follows:573First of all, determine the manufacturing direction of object. Judge whether the mobility ofinsert has to be maintained. These are the key references for selecting suitable manufacturingprinciples. If the insert has to be fixed in the RP object, analyze whether the feature structure ofits shape and the geometric shape would impose any difficulties in times of embedding, so as todetermine whether the support of featured structure should be established or not. The following two points are the factors when considering the positioning of insert:i.

The following two points are the factors when considering the positioning of insert:i. If support of featured structure is not required, a cavity that meets its geometricshape should be designed.ii. If support of featured structure is required, the algorithm of support of geometricfeature structure should be used to design its featured support structure, so as toplace and position the insert and avoid the difficulties and size difference whendesigning the cavity of RP object for the embedding of complex insert. If the insert needs to keep its mobility, we should analyze the feature structure of the external appearance of the insert, so as to judge whether the embedding direction of the insert might cause any interferences. The following two points are the factors for considering themobility of insert:i. If interference occurs, consider changing the forming direction of prototype.ii.

The following two points are the factors for considering themobility of insert:i. If interference occurs, consider changing the forming direction of prototype.ii. If interference does not occur, cover the insert with a thin layer of wax as apre-treatment of the surface of the insert.Upon completion of the above procedures, RP semi-product can be manufacturedimmediately, and then the insert can be embedded. After that, suitable volume of photopolymer should be poured in, and the procedure of vacuum extraction is undertaken. Then theproduct is put into the ultra-violet oven for solidification treatment, acquiring the final product.Finally, the wax layer between the mobile object and the RP object should be removed aftersolidification treatment.

The procedures of online insert embedding process described hereinare shown in Figure 2.2.1 Processes of Online Insert Embedding Approach Focusing on the functional and characteristic differences of insert, the above procedurescan be summarized into three different treatment methods. Detailed description of these threeprocesses is as follows: Process 1：The embedding of a circular shaft with different sectional area, as shown in Figure 3, istaken as an example. In general SLA process, as such work-piece is embedded, the insert has tobe divided into two parts to avoid the problem of recoating interference.

In general SLA process, as such work-piece is embedded, the insert has tobe divided into two parts to avoid the problem of recoating interference. The “online insertembedding process” proposed in this paper adopts the pre-designed depression that meets thefeature structure of the insert, which then embeds the insert according to the required direction.574After embedding, appropriate volume of photo polymer is poured in up to the surface height.Finally, post-cure processing treatment is undertaken to acquire the final product. This processcarries two merits: first of all, the avoidance of the problem of recoating interference, andsecondly, additional dismantling of circular shaft is not required.

This processcarries two merits: first of all, the avoidance of the problem of recoating interference, andsecondly, additional dismantling of circular shaft is not required. Process 2：If the shape of the insert is too complex to design the cavity contour of RP object, whichultimately causes the failure of embedding the insert, then the algorithm of support of geometricfeature structure proposed herein should be employed.While determining the manufacturing direction of RP object, researcher has to considerwhether the shape structure of the insert might cause any problem to the embedding, and undergoan analysis of feature structure. The process is as follows: After horizontal scanning of theshape of the insert, the related information to its shape can be obtained. Then select the feature support curve to establish the RP feature support.

Then select the feature support curve to establish the RP feature support. Next, use normal vector method to judge whether interference will be caused between the established RP feature support of this model and the insert. If there is no such phenomenon,select an appropriate thickness for the growth of solid along the curve, and build a support structure below the solid. If there is interference between the feature support and the insert, thegrowth of solid is not required.Upon completion of the above procedures, input the file into RP system for manufacturingof RP semi-product, and then embed the insert. After completion of embedding, pour inappropriate volume of photo polymer. As the procedure of post-cure processing is undertaken,put the product in ultra-violet oven for solidification treatment so as to obtain the final product.According to the above processes, we can further indicate the various procedures by aschematic diagram in Figure 4.

Process 3：If the mobility of the insert has to be maintained, whether the embedding direction of the insert is workable should be evaluated first. Then analyze the feature structure of the shape ofthe insert so as to judge whether any interferences will occur between the feature structure of theinsert and the RP object. If there is no such interference, proceed with surface treatment of theinsert by pasting a thin layer of wax.Upon completion of the above procedures, a suitable cavity for embedding of the insertshould be designed on the RP object. Next, manufacture the RP semi-product, embed the insert,and then seal the gap between the insert and the RP object by wax in order to avoid the leakageof photo polymer. After the treatment, pour in appropriate volume of photo polymer. Havinggone through procedure of vacuum extraction, the product is put in the ultra-violet oven forsolidification treatment. Then the wax layer is removed to acquire the final product.

Havinggone through procedure of vacuum extraction, the product is put in the ultra-violet oven forsolidification treatment. Then the wax layer is removed to acquire the final product. The575process is shown in Figure 5.2.2 Rules of Lateral Support of Insert In order to support and position the insert effectively, a lateral support of feature algorithmof insert should be developed to collaborate with the online insert embedding process. Throughthe analysis and determination by this algorithm, an RP support can be established along the 3D external contour of the insert. This section describes and investigates into this algorithm.In describing the algorithm, this paper takes the use of SLA for online insert embedding asan example. First of all, a cavity of RP object is designed on the CAD. Next, the RP object ismanufactured, and finally the insert is embedded into the cavity of the RP object.

First of all, a cavity of RP object is designed on the CAD. Next, the RP object ismanufactured, and finally the insert is embedded into the cavity of the RP object. In thisprocess, there should be a tolerance between the RP object and the insert. And this cavity isdesigned according to the deviation of the external structure of the insert to a fixed size tolerance.Since the insert is embedded in the process, the cavity is full of photo polymer. Thus, when the insert is embedded, the residue of photo polymer solvent, which is left in the gap between the cavity and the insert in times of insert embedding, makes it difficult for the insert and the RP object to collaborate with each other, and causing loose in embedding subsequently.As shown in Figure 5, there must be a certain tolerance between the insert and the cavity. The design of a lateral positioning structure at the RP object can improve the matching problembetween the insert and the RP object.

The design of a lateral positioning structure at the RP object can improve the matching problembetween the insert and the RP object. By so doing, the RP object can contact the insert.Besides, position the insert more effectively, a small gap should be left before the processing foreasier control of the pouring of photo polymer.In order to embed and position the insert rapidly and effectively, the lateral support offeature algorithm of the insert is proposed. This paper divides it into boundary projectioncontour, contour curvature definition and establishment of lateral support point, which thenfurther discuss into these three steps.Step 1: Boundary Projection Contour To avoid the creation of interference between the insert and the RP object in the embedding process of the insert, the maximum projection area of the insert in the embeddingdirection should be acquired.

Suppose that when a parallel light irradiates downward from thetop end of the insert, there is a shadow of the insert appeared at the bottom. This shadow is justthe maximum projection contour in such embedding direction. Taking a shaft with changeablesectional area as an example, the maximum projection contour, as shown in Figure 6, can beacquired after projection.The maximum boundary projection contour is calculated by using equations (1) – (2).Suppose that the contour of the projection contour is an enclosed curve composed of nsegments of straight line. Then it can be equally divided into m points in anti-clockwisedirection. The contour boundary S can be defined as:576S = ｛Pj =（xj , yj）︳j = 1,2,3……….m｝ (1)In the above equation, P denotes the coordinates of the contour boundary, and m denotes the number of equation.

The contour boundary S can be defined as:576S = ｛Pj =（xj , yj）︳j = 1,2,3……….m｝ (1)In the above equation, P denotes the coordinates of the contour boundary, and m denotes the number of equation. When S is divided into n subsets, S can be indicated in equation (2) below:Si =｛Pij =（xij , yij）︳j = 1,2,3………m, i = 1,2,3…………n｝ (2)And S can be expressed as:S = S1 S∪ 2 S∪ 3………… S∪ n = S∪ iWhere, Si =｛Pij =（xij , yij）︳j = 1,2,3………m, i = 1,2,3…………n｝Step 2: Contour Curvature DefinitionThe determinant of contour curvature can be calculated according to the definition of Dn expresses with equation (3).r D n=r r TN • n n+ = 1r r TN n n1 +θcos(3)Here Nn denotes the normal vector of point n, Tn+1 denotes the vector of the cutting line n+1, and ● denotes the symbol of dot product, being the angle between Nn and Tn+1.

As shown in Figure 7, this curve is equally divided into m points, in which the determinant of point i is thedot product value of the normal vector of the point i and the cutting line of point i+1. After calculation, it is known that Di < 0, and the segment Si is a convex curve. Similarly, the determinant of point j is the dot product value of the normal vector of the point j and the cutting line of point j+1. After calculation, it is known that Dj = 0, and the segment Sj is a straight line. The determinant of point k is the dot product value for the normal vector of point k and the cutting line of point k+1. After calculation, it is known that Dk > 0, and the segment Sk is a concave curve.

After calculation, it is known that Dk > 0, and the segment Sk is a concave curve. The determinant of contour curvature can determine the concave or convexvariations of the contour, which can be referential for the creation of lateral support.Having acquired the contour of the maximum projection contour in Step 1, and defined thecontour curvature in Step 2, the contour segment combination of the projection contour and theinformation of contour curvature can be achieved. Then, by using Step 3 below, the RP lateralsupport structure can be established for the insert, and the problem of positioning can be solved.Step 3: Establishment of Lateral Support Structure Through the above steps, the contour of the maximum projection contour can be analyzedfrom any insert type. For convenience sake, we can classify the contour of projection contour577into three basic types: line type, curve type and hybrid type.

For convenience sake, we can classify the contour of projection contour577into three basic types: line type, curve type and hybrid type. The establishment of the support point is to find out from the determinant Dn the joint between two segments, divide the contour into several sets of segments, and finally establish lateral support structure at the middle points ofvarious segments. Figures 8, 9 and 10 are the examples of the three basic contour types of the support structure. As the contour is circular, Dn＜0, and there is no variation or joint. Lateral positioning structure should be established at the three equally divided contour shapes, as shownin Figure 11. 2.3 Rules of Bottom Support of InsertBy means of the above-mentioned lateral support, the problems of lateral support and positioning of the insert are solved. Next, the paper discusses the bottom support problem of the insert.

Next, the paper discusses the bottom support problem of the insert. Having analyzed the outline curved surface of the insert, a support structure should be established on the geometric feature at the bottom of the insert. This support structure should reach the weight of the support insert, and achieve the effect of positioning. To establish the bottom support of the insert, there are three steps to be taken: effective visible surface definition, calculation of surface curvature of the insert, and determination of the position of bottom support point.Step 1: Effective Visible Surface Definition In the embedding process of the insert, in order to solve the interference problem betweenthe insert and the RP object effectively, the visible surface in the embedding direction should becalculated.

Suppose that a parallel light irradiates upward from the bottom of the insert, lightshadow will appear on the surface of the insert, which should be the visible side in such direction.Taking a shaft with variable sectional area as an example, the visible surface, as shown in thefigure 12, can be acquired after irradiation, which is known as the effective visible surface.Step 2: Surface Curvature Definition of Insert The aim of calculating “effective visible surface” is to determine the position of bottom support point. Take any random point Pn on the visible surface as the standard point, and find out the dot product of the normal vector Nn of this point (Pn) and the direction vector of the adjacent point (Pi), as shown in Figure 13.The determinant of Cn is specified in the following equation:Cn=r r LN • n n+ = 1r r LN n n1 +θcos(4)From the above equation, the curvature Cn of any random point n on the surface of the insert can be obtained.

As the point corresponds to the k adjacent points, if the determinant Cn is negative, and then this point is regarded as a convex point. If the determinant is positive, this point is regarded as a concave point. But if the determinant is zero, a plane surface will appear between this point and the adjacent point. The equation578Step 3: Determination of Bottom Support Position Through the above definition, an effective visible surface can be analyzed at the bottom of any insert. After that, determinant Cn is used to find out the convex side, concave side and plane surface, and then the position of the bottom support point can be determined.If it is judged to be a convex surface, an upward extending cup-shape support structureshould be established from the lowest curve at the bottom, as shown in Figure 14.

If it isjudged to be a concave surface, a downward extending column-shape support structure should beestablished from the highest point of the depression, as shown in Figure 15. If it is judged to bea plane surface, no support structure is required.By using the “rules of lateral support” and “rules of bottom support” proposed in this paper,the insert can be treated according to the following procedures so as to obtain its supportstructure, and achieve the objective of online manufacturing of insert imbedded. Theprocedures are indicated as follows:1. Decide the embedding direction. 2. Find out the contour of the maximum projection contour. 3. Determine the curvature and the joint of the contour. 4. Construct lateral support structure. 5. Find out the effective visible surface. 6. Determine the curvature of surface. 7. Determine the bottom support point. 8. Construct the bottom support structure.3.

5. Find out the effective visible surface. 6. Determine the curvature of surface. 7. Determine the bottom support point. 8. Construct the bottom support structure.3. Conclusion The online insert embedding process proposed in this paper is to solve the problems currently encountered in the SLA process when manufacturing complex geometric structure, functional mechanical structure and assembling object, such as the problems of laser shadowing, recoating interference, positioning and supporting, etc. The support of feature algorithm is developed in this paper. According to analysis and the future experimental results, online insert embedding process can effectively manufacture assembling object and insert to a satisfactory performance. In the future, the researcher intends to further the study into an in-depth investigation towards the deformity and size precision of insert at the RP object Reference: 1. Bertsch A., Bernhard P. Vogt C.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAPPROACH TO DEFINING THE MAXIMUM FILLER PACKING VOLUME FRACTION IN LASER SINTERING ON THE EXAMPLE OF ALUMINUM-FILLED POLYAMIDE 12A. Tarasova, A. Wegner, G. WittDepartment of Manufacturing Technologies, Institute of Product Engineering, University Duisburg-Essen Duisburg, GermanyAbstract Laser sintering is one of the most popular additive manufacturing techniques that uses thermoplastic polymer powders to generate layer-by-layer complex structures. Despite its broad application, some limitations exist restricting its further development. One such restriction is a narrow assortment of commercially available materials that would allow the production of the parts with the desired mechanical characteristics, which is the case with the widely used Polyamide 12 (PA12).

Reinforcement of a matrix polymer with metal particles is routinely performed to achieve better mechanical properties. In this work, a PA12 system enhanced with a 35% volume ratio of aluminum was investigated. Mechanical characteristics, e.g. elastic and flexural moduli, were examined with respect to variation of manufacturing process parameters. In addition, a new methodology was tested, which should help determine the maximum filler packing volume fraction corresponding to the highest mechanical characteristics of a polymer-filler mixture.KeywordsMaximum filler volume packing fraction, Laser sintering, Polymers, Reinforcement, Fillers, Additive manufacturingIntroduction Additive manufacturing (AM) allows convenient generation of individual and complex structures by layers over a short time compared to conventional methods. This has contributed to the AM method’s broad application in areas such as medicine or airspace [1].

This has contributed to the AM method’s broad application in areas such as medicine or airspace [1]. Laser sintering (LS) is one such technique which has the capacity to develop into the manufacturing technology for small-series production. However, the main obstacle for that is the limited assortment of available materials. Compared to conventional manufacturing (CM), only several thermoplastic polymer powders, predominantly PA12, can be used in LS. As a result, final structures don’t always have the desired mechanical characteristics [2].Reinforcement of a polymer with diverse fillers is widely used to enhance material characteristics, and therefore the properties of the final product [3–5]. The type of filler, its shape and particle size distribution, as well as its material properties, all affect the properties of the final composite. In addition, the amount of the filler introduced is crucial for the mechanical properties [6].

In addition, the amount of the filler introduced is crucial for the mechanical properties [6]. Thus, it is vital to know the maximum amount of the filler that can be introduced to guarantee a favorable reinforcement of the material under examination.Recent studies focused on the investigation of diverse fillers and filling ratios in LS [7– 10]. Karevan et al. [7] investigated PA12 composites with two different mass fractions (wt%) of graphite nanoplatelets (GNP): 3 and 5 wt%. The results were compared with the same composites processed by injection molding (IM). The tests revealed that the best mechanical properties can be977achieved by introducing 3 wt% of GNP. Moreover, the choice of the processing method had almost no influence on the properties. Curling effect, which is known to occur in LS, was studied in detail by Mousa et al [8]. They investigated various PA12 compounds reinforced with 10, 20, 30, and 40 wt% of glass beads.

Curling effect, which is known to occur in LS, was studied in detail by Mousa et al [8]. They investigated various PA12 compounds reinforced with 10, 20, 30, and 40 wt% of glass beads. They showed that the PA12 sample with 10 wt% of the filler demonstrated less tendency towards curling than other mixtures. Yan et al. [9] sintered PA12 filled with 30, 40, and 50 wt% of carbon fibers. They observed that the flexural characteristics improved with increasing the amount of the filler due to the good bonding between the carbon fibers and the PA12. Wudy et al. [10] studied various PA12 / glass beads blends and showed that the bonding to the matrix polymer dramatically decreased with a filler content of 50 vol% resulting in highly porous specimens.In addition, there is no methodology or Rule of Thumb as to how much of the filler should be used in a LS process to guarantee the best properties.

For this reason, estimation of the maximum filler packing fraction (Vf max) is important, because it will ideally allow rational use of resources and a much faster correct approximation of the filler quantity.To the best of authors’ knowledge, there are no methods that describe how the optimal filler content can be calculated for LS, and no research has been performed on this account. In this study, a potential approach to defining Vf max in LS is described and was tested on aluminum-filled PA12 samples. Aluminum was selected as a filler because it effectively enhances mechanical properties of the final composite, is widely used for reinforcement in LS and inexpensive [3, 4]. Moreover, its use guarantees processability of our test compound in LS process.

Moreover, its use guarantees processability of our test compound in LS process. The chosen filler amount and its influence on the mechanical properties of the manufactured specimens were investigated while varying parameter combinations in the LS process.Materials and Methods Considering the excellent processability of commercially available aluminum-filled PA12 in LS and their good finishing properties [3, 4, 11, 12], aluminum powder (R&G Faserverbundwerkstoffe) and PA2201 (EOS) were chosen as the filler and matrix polymer for this study.The maximum filler packing fraction (Vf max) can be obtained from the oil adsorption of a filler (OA) using Equation 1, where ρf is the filler’s density [6, 13]. The oil adsorption of aluminum was determined in conformity with DIN EN ISO 787-5 [14].Vf max=1 1+ ρfOA 100(cid:932)Equation 1It should be mentioned that Vf max depends highly on the packing density (ρpack) of a matrix polymer, i.e.

the higher the ρpack value is, the more filler could be mixed with it [6]. The particle form of the polymer powder and its further processing history define ρpack. Spherical particles represent the ideal case in which they could fill the space reaching the ideal theoretical packing density of 74%. For CM methods this value is typically close to 60%. Polymer powders that are commonly used in LS usually have ρpack in the range of 45 – 50% [2]. Thus, less filler amount is required to successfully reinforce a polymer in LS.Our assumption was that in the ideal case a polymer with the ρpack value of 74% is able to completely accommodate Vf max defined by Equation 1. Remembering that in the case of LS the filler content is lower, we used a direct proportion to describe Vf max (Equation 2).Vf max LS=ρpack∙Vf max 74Equation 2978To test this approach, it was supposed that PA12 has the ρpack of 50%, which is theoretically the maximum value among LS powders.

This value would allow the best possible enhancement of material properties in LS. In practice, LS powders have packing density values lower than 50%. Therefore, a likely decline in the material characteristics should be expected when approaching this value due to excess of the filler [2, 6].After determining Vf max LS for aluminum (35%) according to Equation 2, the filler and the polymer were mixed using a TURBULA Shaker-Mixer (Willy A. Bachofen) for one hour at the nominal speed of 23 rpm. After that the LS processability of the collected mixture was verified by investigating flow and thermal characteristics of the material. The flow characteristics of the powder can be defined by measuring the bulk, and tapped densities, and also by calculating the Hausner ratio [2]. Polymer viscosity was studied by melt volume rate method (MVR) by loading the sample with 5 kg at 235°C, in accordance with DIN EN ISO 1133-1 [15].

Polymer viscosity was studied by melt volume rate method (MVR) by loading the sample with 5 kg at 235°C, in accordance with DIN EN ISO 1133-1 [15]. The processing temperatures were determined by differential scanning calorimetry (DSC) under DIN EN ISO 11357-1 [16] using a Mettler Toledo calorimeter. The mentioned techniques allow to assess the processing capabilities of the mixed PA12 with 35 vol% of aluminum (further referred as PA12- 35Al).The next step was the fabrication of PA12-35Al specimens using a VanGuard HiQ HS machine (3D Systems). During the manufacturing process, various process parameters such as temperature (T), hatch distance (h), laser power (P) were examined and optimized. The layer thickness (d) and the laser speed (Vs) were kept constant at 0.1 mm and 10160 mm/s, respectively. The surface energy density (ES) was calculated using Equation 3 [2]. All conducted experiments are listed in Table 1.Table 1.

The surface energy density (ES) was calculated using Equation 3 [2]. All conducted experiments are listed in Table 1.Table 1. The list of experiments and corresponding LS process parameters of PA12-35Al Laser Power P, W 30.5 40.6 50.8 38.1 50.8 63.5 45.7 61.0 76.2 38.1 50.8 63.5 38.1 50.8 63.5Temperature T, °CHatch distance h, mmSurface Energy Es, J/mm2 0.015 0.020 0.025 0.015 0.020 0.025 0.015 0.020 0.025 0.015 0.020 0.025 0.015 0.020 0.025Experiment1 2 3 4 5 6 7 8 9 10 11 12 13 14 150.201750.250.301780.251820.25979ES=P h∙VsEquation 3After the manufacturing process, the sintered parts were examined with respect to their mechanical characteristics and how they depend from the processing parameters chosen. Density measurements of the solid bodies, tensile and bending tests were then carried out. The density values were calculated using the Archimedes’ method and in accordance with DIN EN ISO 1183- 1 [17].

The density values were calculated using the Archimedes’ method and in accordance with DIN EN ISO 1183- 1 [17]. For the tensile tests, five specimens from each parameter set were sintered in the x-direction. The tensile tests were conducted under DIN EN ISO 527-1 [18] with a testing speed of 5 mm/min using tension bars of Type A (DIN EN ISO 3167 [19]). For the bending tests, rectangular samples (L × W × H: 80 × 10 × 4 mm) were constructed and bent according to DIN EN ISO 178 with a bending speed of 10 mm/min [20].Our experimental results were then compared with those of the commercially available aluminum-filled materials on the PA12 basis: Alumide (EOS) and DuraForm Prox AF+ (3D Systems). This would help us to draw conclusions about the applicability of our approach to defining Vf max LS.Results and DiscussionPowder characteristics First, flow and thermal characteristics of the powder were investigated to assess the processability of the prepared powder mixture.

To evaluate the PA12-35Al powder properties, the pristine PA12 (PA2201, EOS) was used as a reference material and was additionally examined in the same manner as our test powder.Good powder flowability is essential for efficient spreading of the powder across the building platform during the manufacturing process. To evaluate the flowability of our mixture, first the bulk and the tapped densities were determined, and from these the Hausner ratios were calculated (Table 2).Table 2. The bulk and tapped densities as well as Hausner ratios of pure and aluminum-filled PA12MaterialBulk density, g/mlTapped density, g/mlHausner ratio, -PA120.4560.4931.08PA12-35Al0.6520.7391.13It is important to know how good the powder particles fill the space in a defined volume, which can be examined by calculating the specific gravity of the powder under study. For that the densities of the solid bodies for both reference and composite materials have to be estimated.

For that the densities of the solid bodies for both reference and composite materials have to be estimated. The experimentally obtained density of the pristine PA12 (ρm) was ~1.0 g/cm3. The target density of the PA12 mixture with 35 vol% (Vf) of aluminum was calculated using Equation 4 and is 1.6 g/cm3 [5]. The density of aluminum (ρf) was 2.7 g/cm3 as specified by the manufacturer.ρ=Vf∙ρf+(1-Vf)∙ρmEquation 4Therefore, the specific gravities calculated were ~46% for the PA12 and ~41% for the PA12-35Al. Irregular-shaped particles of aluminum are the reason for the observed decreased value of the specific gravity, which results in less volume filling in the PA12-35Al mixture than in the980pristine PA12 powder [21]. As a consequence, more porosity is expected for the sintered parts [2, 21].By comparing bulk and tapped densities, it is possible to acquire information about the powder flowability, and therefore derive the Hausner ratio, as we described above.

In brief, the lower the ratio is, the better the flow characteristics of the powder are. In the case of PA12-35Al, the Hausner ratio slightly increased compared to PA12. Nevertheless, both values are in the range of good powder flowability, which should be under 1.25 according to Schmid et al. [2]. The ratios below this value should theoretically allow smooth powder spreading during the LS process.Good binding between the fabricated layers is vital to achieve superior mechanical properties of the manufactured parts. For this reason, the polymer viscosity was investigated using MVR method. The obtained values for PA12 and PA12-35Al were 73 and 9 cm3/10min, respectively. Proper melt formation is expected if the results are above 30 cm3/10min [2]. The aluminum-filled polymer shows increased viscosity, hence decreased melt flowability.Prior to the manufacturing process, the temperature window in which the processing can be executed has to be estimated.

The crystallization temperature (Tc) and the melting point (Tm), as well as their onset values (Tc (onset) and Tm (onset)) were determined using DSC analysis, with the results presented in Figure 1. Good processability is typically observed between the onset values. The determined processing range of both materials is similar, with the PA12 sample showing the Tc (onset) and Tm (onset) values at 152.2 and 181.9 °C, and the PA12-35Al at 152.2 and 180.9 °C, respectively. Therefore, the processing range of both materials is ~30 K.Tc(onset)Tm(onset)43I=PA12 PA12-35Al(cid:39)T = 28.7 KTm = 186.5 °Cg / W2(cid:39)T = 27.9 K/w o l f1Tm = 185.3 °Ct a e H01Tc = 149.3 °C2Tc = 147.9 °C3100120140 Temperature / °C160180200Figure 1. Crystallization and melting behavior of PA12 and PA12-35AlBased on the acquired results of the flow and thermal characteristics, the PA12-35Al appears processable in the LS.

Crystallization and melting behavior of PA12 and PA12-35AlBased on the acquired results of the flow and thermal characteristics, the PA12-35Al appears processable in the LS. However, the high viscosity might be an obstacle in producing good- quality parts.Laser sintering and the mechanical characteristics of the manufactured parts The PA12-35Al proved to be not difficult to process. The spreading of the powder across the building area was smooth and homogenous, while the melting traces looked flat and slightly sunk in the powder bed. In general, the LS process of the PA12-35Al was similar to that of the981unblended PA12. A minor difference noticed is that the powder spreading was significantly smoother with increasing process temperature in the case of PA12-35Al.The best melt formation, as well as sufficient layer binding, can be expected if the samples show high densities [2], and the higher density is, the better the mechanical properties of the manufactured samples are.

The samples with a hatch distance of 0.15 mm showed some deformation, while a hatch distance of 0.3 mm resulted in smaller density values. Therefore, only a hatch distance of 0.25 mm was further investigated at higher temperatures (178 and 182 °C). We observed that at 182 °C, an energy density of 0.025 J/mm2, and a hatch distance of 0.25 mm the highest density of 1.47 g/cm3 was reached, which constituted ~92% of the target density value (1.6 g/cm3).According to the manufacturing data of EOS and 3D Systems, their aluminum-filled materials based on PA12 have densities of 1.36 and 1.31 g/cm3, respectively, albeit the exact filler content isn’t known [3, 4]. However, the approximate amount of the filler in these materials could be estimated by using Equation 4 and inserting the known densities of aluminum (2.7 g/cm3) and PA12 (1.0 g/cm3). The calculated Vf of aluminum in aluminum-filled materials of EOS and 3D Systems is ~21 and ~18 vol%, respectively.

The calculated Vf of aluminum in aluminum-filled materials of EOS and 3D Systems is ~21 and ~18 vol%, respectively. Therefore, the highest content of aluminum in the tested PA12-35Al can be confirmed.The mechanical properties were examined by performing tensile and bending tests on the parts, which were printed in the x direction.

Flexural strength (a) and flexural modulus (b) of the PA12-35Al samples sinteredaccording to Table 1983The same tendency as with the density measurements was observed here. More specifically, the better mechanical properties were observed with increasing temperatures and energy densities while decreasing hatch distances. For instance, a slight improvement of ~5% in tensile modulus and tensile strength was observed with decreasing hatch distance or increasing process temperature from 175 to 178 °C. By further increasing temperature to 182 °C, the tensile characteristics were found to significantly improve up to 40% in comparison to the samples manufactured at 175 °C. Among all specimens the elongation at break was approximately 1.5%, which is rather too low for the thermoplastic polymers and points to the rigidness and brittle behavior of the manufactured samples [22].

However, such behaviour is common among the reinforced polymers which have significant losses in elongation at break due to the introduction of a filler [22, 23].With decreasing hatch distance both flexural strength and flexural modulus increased up to 20 and 15%, respectively. The samples sintered at 182 °C demonstrated a further enhancement in flexural properties up to 25%.The maximum tensile and flexural moduli of the specimens were observed at an energy density of 0.025 J/mm2, a hatch distance of 0.25 mm, and at 182 °C. The maximum values of tensile and flexural moduli were 4210 and 3850 MPa, respectively. These values are ~60% higher than that of the samples processed at a lower energy density of 0.015 J/mm2 and at 175 °C.The commercially available aluminum-filled materials from EOS and 3D Systems demonstrate slightly improved flexural and tensile characteristics (Table 3), though the filler content is supposed to be lower than in the investigated PA12-35Al.

This means that anticipated enhancement of mechanical properties didn’t occur. On the one hand, the reason could be the low adhesion between PA12 and aluminum powder since the latter wasn’t surface-treated which guarantees superior bonding between filler and polymer particles [6]. On the other hand, the chosen filler quantity introduced into the polymer matrix could be too high, therefore resulting in less than optimal coverage of the filler by the polymer [6].Table 3. Mechanical characteristics of commercially available aluminum-filled PA12 [3, 4]MaterialTensile Modulus, MPaTensile Strength, MPaElongation at break, %Flexural Modulus, MPaFlexural Strength, MPaAlumide (EOS) DuraForm Prox AF+ (3D Systems)38004340483743360037107264Conclusions The new approach of defining the maximum filler packing fraction in the polymer system for the laser sintering was tested.

Due to good processability of the commercial aluminum-filled powders, PA12 and aluminum were chosen as a matrix polymer and a filler for the test compound. Of importance was to verify the applicability of the proposed approach by examining mechanical properties of sintered specimens. The results of the performed mechanical tests showed that the tested filler amount of 35 vol% proved to be above the limit where no enhancement in properties can be expected. This can be also observed by comparing mechanical characteristics of PA12- 35AL with aluminum-filled materials from EOS and 3D Systems, both of which supposedly have less aluminum incorporated in the PA12. The experimentally determined density values of PA12- 35Al were ~8% lower than the target value of 1.6 g/cm3, which might indicate the presence of some984aluminum particles that are not completely bonded to the melted polymer, therefore resulting in porous and brittle structures.

As a result, no enhancement was observed in the tested samples.The observed decline of the mechanical properties was taken into consideration by assuming the initial packing density of PA12 with a maximum value of 50%. Therefore, the proposed and tested approach based on estimating the optimal filler volume packing fraction for laser sintering can be considered viable. However, a further testing of different filler ratios is needed in order to correlate them with the actual packing density of a polymer.Acknowledgements The presented study was supported in the IGF project 19646 N by the Federal Ministry ofEconomics and Energy based on the resolution of the German Bundestag.References[1] T. Wohlers, Wohlers Report 2017 - 3D Printing and Additive Manufacturing State of theIndustry. Fort Collins: Wohlers Associates Inc., 2017.[2] M. Schmid, Laser Sintering with Plastics: Technology, Processes, and Materials. München:Carl Hanser Verlag, 2018.[3] EOS GmbH.

PROACH TOLERANCE IN THE ASSEMBLIES OF EVOLUTIONARY HYBRID PROTOTYPESM. Rivette, P. Mognol and JY. Hascoet Institut Recherche Communications Cybernetique Nantes (IRCCyN) UMR CNRS 6597, 1 rue de la noe, BP92101, 44321 Nantes Cedex FranceReviewed, accepted September 14, 2006Abstract A new answer is proposed to replace the traditional “one shot” prototype (manufactured in one piece with one process): the hybrid rapid prototype. It is used to highly reduce time, cost and increase reactivity during the development times of new products. The part is decomposed in several components which can quickly be changed and can be manufactured with a process the most adapted. The main objective of the presented method is to propose an available technological assembly between the different components of the part in the respect of technological and topological function, and initial tolerance.

Using a graph of representation, fuzzy logic and a tolerance point of view, some entities are associated with a CIA (Assembly Identity Card) in accordance with evolutionary and manufacturing analysis. This work will be illustrated by an industrial tooling for plastic injection.1 INTRODUCTION1.1 Context Modern means for the improvement of efficiency in manufacturing are quite diverse. However, the traditional cornerstones, cost, quality and time are the targets that business is managed. Thus, mass customization is one of the modern means to achieve these goals, and allows customizing product to individual clients and producing those with principles of mass production. For many products, the competitive edge of the producer is dependent not merely on the price, but also on the choices or variations provided in each product line.

For many products, the competitive edge of the producer is dependent not merely on the price, but also on the choices or variations provided in each product line. Examples of such products range from automobiles (manifested by the increasing number of "variants" available in any base model of a car) to electronic products such as computers. The challenge is to create a variety of products from a common family without a significant trade-off in production costs or lead time. In order to address the high cost of this practice, manufacturers develop product families from a common platform that is shared by all the products, whose variants are designed to fulfill different customer demands. The variants are created by adding specific components to the platform. The use of a small batch with an alternative is thus an increasingly frequent option. At the same time, products have become more complex and the development times of products are increasingly long and laborious.

At the same time, products have become more complex and the development times of products are increasingly long and laborious. In this context, the hybrid rapid prototype is appearing. In this paper the rapid tooling and the prototype part are presented. The tools can be of evolutionary, for bridge tooling or for small series. The parts can be aspects, geometrical, functional or technological prototypes.5731.2 Related Work Hybrid rapid prototyping has been presented in some recent papers [1, 2]. Their researches were based on the study of two processes: CNC and Rapid Prototyping (RP) systems. CNC is used when the quality of the part is higher than what it is possible to obtain with RP system. Their methodology uses STEP AP-203 data but didn’t take into account ISO specifications of the model. Likewise, concerning the manufacture of the parts, methodology that would allow the decomposition of a part into a space partitioning, was developed [3].

Likewise, concerning the manufacture of the parts, methodology that would allow the decomposition of a part into a space partitioning, was developed [3]. The weakness of these studies is that only build accessibility is considered. A choice of an adequate process is not proposed. At the same time, the assembly of the hybrid prototypes and its repercussions on the design was never taken into account. These works do not envisage any interaction between manufacture, assembly and design. Some research about product families [4, 5], and design for assembly [6]] related only the assembly of subset of a product (for example: a power supply). This work does not treat assemblies of cast solid product composed of hybrid elements (for example: injection tooling). In the field of tool-maker, tools are usually made in several components. This decomposition is the result of knowledge of the engineer but there is no reliable method to obtain this decomposition.

In the field of tool-maker, tools are usually made in several components. This decomposition is the result of knowledge of the engineer but there is no reliable method to obtain this decomposition. In this paper, a novelty method is proposed about the design of this kind of tools. The concept proposed in this paper aims to decompose the part on a Multi Component Prototype (MCP), instead of a part made in one piece. The two main reasons are to include the evolutionary requirement of the prototype regarding to the tests that are performed on it; and to optimize the manufacturing process locally, regarding to the component geometry and functional requirements. Therefore, only one or few components would have to be remanufactured separately, in order to update prototype geometry for testing purposes. A methodology is proposed for assemble the multi component part by using the extraction and the use of entities from a CAD model.

A methodology is proposed for assemble the multi component part by using the extraction and the use of entities from a CAD model. A new part decomposition for a prototype is proposed in order to guarantee the functionality requirements and to allow the evolutionary of its geometry. Furthermore each component of the new partitioned part is built with the more appropriate process. The assembly of components is design to have the same tested characteristic as the “one piece” part. This new approach is entitled “hybrid evolutionary prototypes”. These different analyses are based on CAD STEP specifications and more especially on Application Protocol AP-224. STEP AP-224 is a manufacturing features oriented description. This chapter of ISO 10303 specifies the information needed to define product data necessary for manufacturing a mechanical part. The product data is based on existing part designs that have their shapes represented by machining features.

The product data is based on existing part designs that have their shapes represented by machining features. Chapter AP224 contains all of the information and capabilities necessary to manufacture the required part.2 MCP CONCEPT2.1 Presentation MCP concept is presented in the Figure 1. The goals are to allow evolution of parts for testing purpose. All developed methodology is based on feature analysis. For that, automatic feature recognition [7] like in various research works made in CAPP [8], was not used. Step-AP224 entities in the CAD model of the part is used, because it is essential to574575analysis [10] proposes the best manufacturing process for each component of the studied MCP.

Step-AP224 entities in the CAD model of the part is used, because it is essential to574575analysis [10] proposes the best manufacturing process for each component of the studied MCP. Hereafter, the final analyses are developed.CAD Model AP-224 entitiesManufacturing analysisFeasability analysisEFCCreation of the graph of representationResolution of the grapheFuzzy analysis for manufacturingFCFuzzy analysis for assembly (CIA Concept)MCPFigure 2: The different activities for MCP concept.2.3 Method In the methodology, the F.C concept, presented in chapter 2.1 is applied. A FC is a real component of the hybrid rapid prototype. It is composed with AP-224 feature that participate at the same function regarding to the ISO tolerances (qualitative) and topologic information (Figure 2). The method is presented on an industrial example. For this product, there are two versions, the version number 1 and the number 2 (Figure 3 and 4).

The method is presented on an industrial example. For this product, there are two versions, the version number 1 and the number 2 (Figure 3 and 4). Currently, there are two moulds to manufacture the two version of product. The starting point is Ap-224 entities of the CAD model (Figure 5). This extraction of entities from Ap-224 programs is possible. For example, the commercial software STEP-trans [11] used for obtaining this entities from a CAD model, gives automatically the list of the features Ap-224 on a “simple part”.576Figure 3: Industrial example (Version 1 & 2).Figure 4: 2D draw with functional requirements only for the entity 6.Figure 5: Entities AP-224 mould.The method is used to obtain a hybrid mould allowing: - A realization with the best processes. - The manufacture of the two alternatives products.Manufacturing analysisDuring the manufacturing analysis, each AP-224 entities is analyzed to define the best manufacturing processes to manufacture them.

For each studied processes, each entities received marks (1 to 10). Mark 0, if the process is not available to mark 10 if the process is perfectly adaptated to manufacture the entity. [12] To obtain these marks, 4 processor are used. The first treatment occurs with the feasibility processor. It will estimate the feasibility degree of each feature done by a process. The cost processor estimate the manufacturing cost of each feature, without considering “fixed costs”. The time processor estimate the manufacturing time of each feature, without considering non-productive times. At the end, Fixed costs and non-productive times are taken into account by the classification processor. This processor gives all modules (features gathered together) with their assigned process, and the global estimated manufacturing time and cost. Here for example (Figure 6), the entity 5, injection point with conical forms, will be preferably manufactured with EDM Process.

Here for example (Figure 6), the entity 5, injection point with conical forms, will be preferably manufactured with EDM Process. Nevertheless, the DMLS process can be used but the HSM process is completely unsuited.577Manufacturing capability coefficientEntities HSMEDM DMLS0 1 2 3 4 5 6 7 8 910 10 10 10 10 0 10 10 10 101 5 5 5 5 10 3 5 5 51 1 1 1 1 5 7 3 3 3Figure 6: Manufacturing capability coefficient.Feasibility analysisThe feasibility analysis analyses functional (ISO specifications), evolutionary criterions and topologic configuration between AP-224 entities of the part. The goal is to group together features that participate to the same functions in the same piece of the puzzle. It is considered that features that are highly positioned from a qualitative point of view or which have particular topologic configuration, must be in the same FC.

It is considered that features that are highly positioned from a qualitative point of view or which have particular topologic configuration, must be in the same FC. In the example the feasibility analysis, give the AP-224 entities list and their properties, the list and the details of the links between AP-224 entities (see figure 7 a,b). entity 0 1 2 3 4 5 6 7 8 9TypeToleranceInclusion Contact Intersection(cid:1318) (cid:600) (cid:320)L1 L2 L3 L4 L5 L6 L7 L8 L9L1 L2 L3 L4 L5 L6 L7 L8 L9 0(cid:1318) (cid:1318)Low Low Low Low High Low High High High(cid:1318) (cid:1318) (cid:1318) (cid:600) (cid:600) (cid:600) (cid:600)123 (a)456789(b)Figure 7: List of the links (a) and Details of the Links (b).Here for example (Figure 7 a,b), there is the link number 5 between the entity 5 and the Base shape. The type of this link is inclusion. That means that the entity 5 is including in the Base shape.

The type of this link is inclusion. That means that the entity 5 is including in the Base shape. (The volume of entity 5 is completely contained in the volume of the base shape). Following theses stages, a graph of representation is made in accordance with the rules previously defined (Table 1). This graph of representation proposes a cutting of the part in MCP according to the results of the analysis of feasibility (Figure 8).57843215(cid:938)(cid:938)(cid:938)(cid:938)(cid:938)BASE_SHAPE(cid:600) Contact (cid:1318) Inclusion (cid:136) Intersection(cid:600)6(cid:600) 7(cid:600) 89(cid:600)Figure 8: Graph of representation of the example.A graph of representationThe entities AP-224 extraction, special specification, the functionality constraint product, the manufacturing and feasibility analysis give some important information.

The resolution of the graph representation is used to gather the FC in the respect of the “functionality” of the part. This resolution is based on few rules (table 2).5791Resolution Rules : 1/ The FC containing an evolutionary entity must be smallest as possible (i.e. to contain less possible entity) 2/ A regrouping does not cut a double Line (Unbreakable) 3/ The MCP must be compose at least of Possible FC. PossibleTable 2: Resolution Rules.If possible, the evolutionary FC doesn’t be gathered. In fact, it will be changed during the tests or during alternative production. The objective is to have a minimum number of components for the Hybrid rapid prototype. In this stage the information contained in the link between AP-224 Entities are used. For the example, the graph of representation (figure 8) is obtained.Fuzzy logic analysis for manufacturing.A manufacturing fuzzy logic method is used to find the processes the most adapted to manufacturing FC or FC groups.

During the manufacturing analysis, made before, a feasibility processor estimates the feasibility degree of each feature done by a process. Thus, each feature obtained marks during this analysis. The manufacturing fuzzy logic method use this marks. The objective is to manufacture the complete FC or FC groups with the same process, or if it is not possible to have an available combination of processes. This combination is evaluated by fuzzy logic processor. In the example (Figure 9.) the entities of FC1 are preferably manufactured with HSM process expect the entities 5 witch is preferably manufactured with EDM. In this case the two processes are used successively. The manufacturing analysis proposes the best manufacturing process for each component of the studied MCP.

These three fuzzy logic methods are used simultaneously. At the end of the iterations a solution of gatherable Hybrid rapid prototype is proposed. To manufacture the various alternatives of the product, the CIA 1 is replaced by the CIA2. (Figure 10)Figure 10: Proposed solution.In the following chapter the tolerance point of view of the CIA assembly is developed.3 ASSEMBLY DESIGN AND TOLERANCE POINT OF VIEW3.1 Introduction The main objective of this analysis is to propose an available technological assembly between the different components of MCP. For this, standard assemblies (noted CIA (assembly identity card)) have been defined. Each standard assembly is parameterized by its CIA. It is necessary to associate via fuzzy logic, for each entity a completely definite assembly in order to obtain technological solution of MCP.3.2 CIA Concept A CIA is an identity card of one assembly, which gathers the general characteristics of this assembly.

Each CIA has several parameters, witch completely defines its geometry. The CIA is an informational tool to define physically the assembly between the functional components (FC) of the prototype. The quality of the interfaces between parts ensures the assembly requirements of a mechanism and the right positioning of functional surfaces. During functional tolerancing of an industrial mechanism, designers define the operating mode of the mechanism and impose functional requirements. To formalize the design intents clearly for each CIA, a method, named TLIC (Tolerancing in Localization with Influence of the Contacts) [13], is used. This method usually used in mechanics is applied to the field of Rapid Prototyping. This method uses positioning tables of the parts to clearly indicate the set up surfaces581associated as features and the preponderance order of the features.

This method uses positioning tables of the parts to clearly indicate the set up surfaces581associated as features and the preponderance order of the features. Algorithms of TLIC [14, 15] method generate tolerancing of surfaces of junction between parts and give, for each requirement, the loop of contacts, the active parts and the corresponding inequation. The synthesis of tolerances uses fuzzy expression of requirements. (Figure 3 and 4). From a 2D draw (Figure 4), the fuzzy logic analysis for assembly (the CIA fuzzy logic) chooses in a list of the parameterized CIA, the most adapted CIA for integration of the FC or FC group. In this data base, each CIA has tolerances to respect the functional CIA requirements (figure 11 and 12). After this choice, the TLIC method [16] is used to take into account the functional product requirements. (figure 4). Then the CIA is completely define. (Figure 13).Figure 11: TLIC Method.

This CIA, must take into account the functional product requirement defined by the designer during the conception. The fuzzy logic approach compares the design drawing requirements with the CIA capability informed in the Assembly identity card (Figure 12).This method with tolerance point of view gives some interesting solutions. The TLIC method, usually used in mechanism, is used to obtain the tolerance of CIA in the field of Rapid Prototyping. For this, some adaptations are made on the original TLIC method. First of all, each CIA has dedicated tolerancing. After, in accordance with functional specifications of the part, each assembly between the different components is calculated with a specific method developed by the team. Finally, the numerical value of tolerance for complete assembly (Base_shape and components) is obtained.4 CONCLUSIONThis method takes place during the product development.

Finally, the numerical value of tolerance for complete assembly (Base_shape and components) is obtained.4 CONCLUSIONThis method takes place during the product development. It allows to reduce consequently time and cost when complex prototype must be manufactured and optimized. This methodology uses multiple point of view and knowledge to analyze prototypes definitions. The assembly and geometrical analysis based on 2 different Fuzzy logic583processes, give FC groups. These groups are associated with a CIA in accordance with evolutionary and manufacturing analysis. The TLIC Method is directly used for tolerancing the CIA. The Hybrid Rapid Tools are composed by the assembly of different CIA with the base shape.This work was carried out within the context of the working group Manufacturing 21 which gathers 11 French research laboratories.

The topics approached are: the modeling of the manufacturing process, the virtual machining, the emerging of new manufacturing methods.This work was supported by the Brittany District.5 REFERENCESHur, J., Lee, K., Zhu-hu, and Kim, J., Hybrid rapid prototyping system using machining and deposition. Computer-Aided Design, 2002 34(10), 741-754 Chen, Y.H. and Song, Y., The development of a layer based machining system. Computer-Aided Design, 2001 33(4), 331-342 Ki, D. and Lee, K., Part decomposition for die pattern machining. Journal of Materials Processing Technology, 2002 130-131, 599-607 Erens, F. and Verhulst, K., Architectures for product families. Computers in Industry, 1997 33(2-3), 165-178 Jiao, J., Tseng, M.M., Duffy, V.G., and Lin, F., Product family modeling for mass customization. Computers & Industrial Engineering, 1998 35(3-4), 495-498 Barnes, C.J., Jared, G.E.M., and Swift, K.G., Decision support for sequence generation in an assembly oriented design environment.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAPPROACHES TO GEOMETRIC DATA ANALYSIS ON BIG AREA ADDITIVELY MANUFACTURED (BAAM) PARTS G. D. Dreifus*, *, Y. Jin†, N. Ally*, B. K. Post*Manufacturing Demonstration Facility, Oak Ridge National Laboratory NTRC II, 2370 Cherahala Blvd Knoxville, TN, 37932† Mork Family Department of Chemical Engineering and Materials Science, University of Southern California, Los Angeles, CA, 90007AbstractThe promise of additive manufacturing is that a user can design and print complex geometries that are very difficult, if not impossible, to machine. The capabilities of 3D printing are restricted by a number of factors, including properties of the build material, time constraints, and geometric design restrictions.

The capabilities of 3D printing are restricted by a number of factors, including properties of the build material, time constraints, and geometric design restrictions. In this paper, a thorough accounting and study of the geometric restrictions that exist in the current iteration of additive manufacturing (AM) fused deposition modeling (FDM) technologies on a large scale are discussed. Offline and online methodologies for collecting data sets for qualitative analysis of large scale AM, in particular Oak Ridge National Laboratory’s (ORNL) big area additive manufacturing (BAAM) system, are summarized. In doing so, a survey of tools for designers and software developers is provided. In particular, strategies in which geometric data can be used as training sets for smarter AM technologies in the future are explained.IntroductionAdditive manufacturing (AM) is a $4.1 billion industry that is currently growing at a rate of 35.2% [1].

AM is a means of depositing material layer-by-layer to build a part from the bottom up; in fused deposition modeling (FDM), an extruder-based printer deposits a bead of material along a tool path onto a printer bed [3]. Oak Ridge National Laboratory’s (ORNL) Manufacturing Demonstration Facility (MDF), along with Cincinnati Inc (CI)., designed a large scale 3D printer capable of depositing material at rates almost 100 times that of typical desktop printers [7, 8]. In large scale AM, successful prints are determined by a number of varying factors. The speed of the printer gantry and the bead width of the printed material can fluctuate unexpectedly and so represent noise extent throughout the printing the process. Also, the materials and thermal properties inherent in the print create noise during the process.

Also, the materials and thermal properties inherent in the print create noise during the process. As a particular material cools, warping often occurs in unpredictable ways, propagating error throughout a printed part.The geometry of a given print affects all of these parameters. A part with high curvature or sharp, sudden angles in its geometry will cause high acceleration on the printer head motion, and acceleration creates sudden changes in the flow rate of a printed material. Furthermore, while a proportional-integral-derivative (PID) controller might be used to stabilize the velocity of a printer head, a feedback mechanism to keep the acceleration stable would operate too slowly for the system to remain stable. Thus, high curvature and sharp angles create fluctuations in the bead1922width during a print.

Thus, high curvature and sharp angles create fluctuations in the bead1922width during a print. Geometry can also affect a print as if is highly variable, as this will likewise cause changes in the printer head velocity.This noise exists on the perimeter and interior of the print. On the perimeter, noise is cause by changes in velocity. In the interior, changes in velocity are similarly inevitable, but the infill tool path is influenced by the geometry of the part too. Tool path density and trajectory will have an impact on print viability and build success.To combat these obstacles, scan data on rudimentary part geometries was collected to begin development of an open source training set that might be used to better learn tool paths and do analysis on AM build failures. The literature on deviation analysis in BAAM printing is seemingly sparse. Work has been conducted to understand deviation in stereolithography [10,11].

The literature on deviation analysis in BAAM printing is seemingly sparse. Work has been conducted to understand deviation in stereolithography [10,11]. Deviation models have been developed that consider parametric models of FDM printing that seeks to correct for geometric part error in the CAD model [5,12]. A deviation analysis is presented here for large scale systems and potential strategies to overcome part defects in a given geometry are consequently discussed. This part serves as inspiration for potential intersections of machine learning, tool path planning, and controls.ExperimentsA series of triangles with a variety of angles on the BAAM-CI system was printed. The BAAM- CI printer was developed by ORNL in conjunction with Cincinnati Inc. and can print components as large as 20-ft. long, 8-ft. wide, and 6-ft. tall. It uses a single-screw extruder to deposit feedstock material [7]. For this experiment, 20% carbon fiber reinforced ABS was used.

long, 8-ft. wide, and 6-ft. tall. It uses a single-screw extruder to deposit feedstock material [7]. For this experiment, 20% carbon fiber reinforced ABS was used. The pellet feedstock material is ~3.45mm in diameter and ~3.75mm in length. Triangles were , and 90 printed whose CAD models had angle ratios of 90 50 and holes of various sizes in it was also printed, as can be seen in a topographical deviation analysis image in table 7. A sinusoidal wave with 6-in. amplitude appended to a 43.99-in. by 3.99-in. block was printed as well and can be seen in table 7.60-80 , as can be seen in figures 1-2 and 4-7. A 12-in. by 42-in.

A sinusoidal wave with 6-in. amplitude appended to a 43.99-in. by 3.99-in. block was printed as well and can be seen in table 7.60-80 , as can be seen in figures 1-2 and 4-7. A 12-in. by 42-in. rectangular block with pillars ° ° , 901030702040°°°°°°°°°°Table 1: The dimensions of the triangle prints completed on the FaroArm Laser Scanner.BAAM Prints 80-10 2.12’’ Base Length (in.) Height (in.) Hypotenuse Length (in.)70-20 4.37’’12’’ 12.19’’ 12.77’’12’’60-30 6.93’’12’’ 13.86’’40-50 10.07’’12’’ 15.66’’Sinusoidal 43.99’’9.99’’ N/AScanning AM parts to test deviation analysis and understand part quality has been conducted [4]. To build offline training sets, a Faro Platinum Laser Arm Scanner was used to scan point clouds of each object. Mesh triangulations of a surface geometry were calculated using 3D Systems’ 2015 Geomagic Software, yielding a deviation analysis between the scanned point clouds and the CAD models.

Mesh triangulations of a surface geometry were calculated using 3D Systems’ 2015 Geomagic Software, yielding a deviation analysis between the scanned point clouds and the CAD models. The training sets can be seen in tables 1 through 7. Each table lists a range, denoted “>=Min” to “<Max,” of deviation in the orthonormal direction from the surface192340-50 42’’12’’ N/Ageometry of the CAD model and the number and percentage of points within the point cloud that exist within that given range.1924Previously, infrared camera footage was used to analyze AM metal parts [2]. For online experiments, a FLIR Laser Scanner was attached to ORNL’s BAAM system. The BAAM system was developed by ORNL and is capable of printing components 3.28-ft long, 7.87-ft wide, and 5.9-ft tall. It runs a single-screw extruder using pellet feedstock material, in this case 20% carbon fiber reinforced ABS [7]. The pellet feedstock size is ~ 0.136in. in diameter and ~0.148in.

It runs a single-screw extruder using pellet feedstock material, in this case 20% carbon fiber reinforced ABS [7]. The pellet feedstock size is ~ 0.136in. in diameter and ~0.148in. in length.Experimentation was conducted on the ORNL BAAM machine with an infrared FLIR AX5 camera. The camera has a focal length of 19-in. so a clamp was designed to attach the FLIR camera to the side of the extruder at the requisite distance from the printer bed. The camera runs at a frequency of 30 Hz and has a resolution of 320 by 256 pixels. An online bead width measurement of a given print can now be captured, giving an accurate online measurement of printer flow rate.The FLIR camera can capture online measurements of the width of the bead extruded from the printer head. By clamping a camera to the side of the extruder, flow rate measurements can be read throughout a given print.

By clamping a camera to the side of the extruder, flow rate measurements can be read throughout a given print. While FLIR cameras capture temperature alone, we used a MATLAB filter to derive the width of the bead from differences in the temperature values captured in the pixels. The color values can be captured in a matrix and filtered into a binary grey scale value.The BAAM system’s controls run off the machining language G-Code which assigns the location and velocity of the gantry that moves the extruder. We can run the MATLAB filter on a video feed of the print. Using this and the video feed from the FLIR camera allows for the calculation of a local deviation analysis of the printed material for a given part.

Since acute angles are a frequent cause of print error, the angle width in a layer of a part can be identified within a scan and added to the database that currently includes a few triangles. Error likelihood can be estimated as the training set grows. Similar understandings can be developed for areas with large numerical curvature and the shapes seen in tables 6 and 7. Correlation can be calculated between subset geometry size and curvature.Similarly, analysis can be achieved through an online feedback mechanism. By attaching an IR camera to large-scale 3D printer extruders, deviation analysis can be conducted locally on the interior of part geometries. Statistical correlation can be calculated between numerous parameters of AM, including gantry velocity, flow rate, and tool path curvature.

Statistical correlation can be calculated between numerous parameters of AM, including gantry velocity, flow rate, and tool path curvature. Infill tool path geometry can be classified in much the same way that the surface analysis can be studied using offline scans.Training sets of these parameters, in both the online and offline setting, can lead to fairly basic but extremely helpful tools for designers and software developers. The mean deviation associated with various levels of curvature of a part can be easily computed. Computing the eigenvalues and eigenvectors of the correlation matrix of the surface deviations of various curvatures and knowing consistently successful geometries can help predict the likelihood of success of a print, as large eigenvalues indicate large variance from successful prints.

Further data analysis and machine learning technologies can be applied to calculate how geometric variety affects the success of a print.Similarly, in the online case, the likelihood of the success of a particular tool path can be classified by local curvature. If a particular segment of tool path exists frequently in poor-quality prints, as determined by the offline scan data, then design software can flag a user that a part geometry is not amenable to a print’s success. If certain curvature values can be associated with successful prints, then a classification of tool paths can be constructed, whereby tool path optimization can be achieved. If outlier curvature values in a typically successful tool path frequently leads to print failure, then software can potentially automate the rejection of designs or tool paths containing those trajectories.

These are all examples of potential strategies for utilization of the data acquired via the training sets developed.The data from our tables show where particular part error occurs. As expected, sharp turns result in large protrusions around acute angles. For some prints, we get very narrow distributions, with most deviations extant around only a few standard deviations within each table. For instance, for the 80-10 triangles, 55.64% of the points were within 0-.1 cm from the CAD model; for the 50-1928�40 print, 38.39% of the points were with 0.0044 and 0.1036 cm of the CAD model. For other prints, we get more spread out distributions. Greater sample sets would needed to determine a trend.An analysis of the data begins to demonstration when part failures and defections are more likely. At highly acute angles, we can see part defects and missing data from the original CAD model. We can also see more frequent and sever gap regions extent in triangles with greater acute angles.

At highly acute angles, we can see part defects and missing data from the original CAD model. We can also see more frequent and sever gap regions extent in triangles with greater acute angles. This occurs because, on the scale at which BAAM takes place, higher acceleration causes greater variance in bead width, propagating part defects throughout a print. This can also be seen in the pillars and holes build, in which defects can be found around the pillars and holes. In this print in particular, we learn that there is a limit on “pillar” size in a print. For the three smallest pillars in the print, the build experiences underfill problems, whereas the other pillars resulted in overfills. We also see missing data in the highest curvature regions of the sinusoidal prints.ConclusionIn large scale AM systems, large data sets are needed to better understand the technology.

We also see missing data in the highest curvature regions of the sinusoidal prints.ConclusionIn large scale AM systems, large data sets are needed to better understand the technology. By accumulating training sets on print quality, analysis can be done to correlate particular design parameters, tool paths, and printer settings with a successful build. The acquisition of such data can ultimately be used to program smarter robotic systems for better AM technologies.To achieve these ends, two ways to acquire geometric deviation analysis on ORNL’s BAAM systems were outlined. A series of sample shapes were printed and laser scanned. Using the point clouds acquired from the laser scan, the difference between the cloud data and the surface of a CAD design was tested. This gave an online database which can now be drawn on for various purposes in software design. An online feedback system using an IR camera was also set up on the BAAM system.

This gave an online database which can now be drawn on for various purposes in software design. An online feedback system using an IR camera was also set up on the BAAM system. By collecting a video stream of the print, a deviation analysis can be similarly conducted and a training set built up, this time at each instance of the print.With methods of developing training sets in hand, quality testing on every single printed object of the BAAM system can begin. These training sets can also be utilized in machine learning algorithm development. In this way, BAAM machines can be “taught” when prints will be successful, to what degree, and how they should react to improve print quality and success likelihood.Acknowledgment of FundingThis material is based upon work supported by the U.S.

Department of Energy, Office of Science, Office of [insert the sponsoring SC Program Office, e.g., Basic Energy Sciences], [Add any additional acknowledgments or information requested by the sponsoring SC Program Office] under contract number DE-AC05-00OR22725.1929Notice of CopyrightThis manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05- 00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, worldwide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes.

Elliot, B. Post, R. Smith, and C. Blue, "The Importance of Carbon Fiber to Polymer Additive Manufacturing." Journal of Materials Research, 29(17), pp 1893-1898 (September 2014).[9] B.V. Ramana. Higher Engineering Mathematics. Tata McGraw-Hill Publishing Company Limited, 2006. New Delhi.[10] Reeves, Philip E. (1998) Reducing the surface deviation of stereolithography components. PhD thesis, University of Nottingham.1930[11] Reeves, P.E. and Cobb, R.C., “Surface deviation modelling of LMT processes ‐ a comparative analysis”, 5th European Conference on Rapid Prototyping & Manufacturing, Dipoli Conference Centre, Helsinki, Finland, 5‐7 June 1996.[12] Xu, L., Huang, Q., Sabbaghi, A., and Dasgupta, T., 2013, “Shape Deviation Modeling for Dimensional Quality Control in Additive Manufacturing,” ASME Paper No. IMECE2013- 66329.1931.

ueous-Based Extrusion Fabrication of Ceramics on DemandMichael S. Mason1, Tieshu Huang1, Robert G. Landers1, Ming C. Leu1, Gregory E. Hilmas2, and Michael W. Hayes31870 Miner Circle Department of Mechanical and Aerospace Engineering1 Department of Materials Science and Engineering2 University of Missouri-Rolla, Rolla, Missouri 65409-0050 The Boeing Company3, St. Louis, Missouri {mmason,hts,landersr,mleu,ghilmas}@umr.edu michael.w.hayes2@boeing.comReviewed, accepted September 4, 2007ABSTRACTAqueous-Based Extrusion Fabrication is an additive manufacturing technique that extrudes ceramic slurries of high solids loading layer by layer for part fabrication. The material reservoir in a previously developed system has been modified to allow for starting and stopping of the extrusion process on demand. Design pros and cons are examined and a comparison between two material reservoir designs is made.

Design pros and cons are examined and a comparison between two material reservoir designs is made. Tests are conducted to determine the optimal deposition parameters for starting and stopping the extrudate on demand. The collected test data is used for the development of a deposition strategy that improves material deposition consistency, including reduced material buildup at sharp corners. Example parts are fabricated using the deposition strategy and hardware design.INTRODUCTIONAqueous based extrusion fabrication (ABEF) of ceramics is a Solid Freeform Fabrication (SFF) process that deposits ceramic paste layer by layer through extrusion. The process has been previously introduced, and details regarding the build material and equipment can be found in [1-3]. A process schematic and photograph of the system are shown in Figures 1 and 2, respectively. The material deposition device previously developed gave good results for continuous material deposition.

A process schematic and photograph of the system are shown in Figures 1 and 2, respectively. The material deposition device previously developed gave good results for continuous material deposition. The ability to start and stop the material extrusion on demand was not yet developed in our previous work. With the need to be able to fabricate more complex geometries, different extrusion processes were surveyed for the process of starting/stopping material deposition on demand. The first system surveyed was the Fused Deposition of Ceramics (FDC) process, which modified a Fused Deposition Modeling (FDM) machine to deposit a ceramic/polymer fiber strand [4-5]. The FDM process uses two rollers with heaters to move the material fiber either forward for deposition or in reverse to stop the deposition. The second process uses a Micro-Scale Robotic Deposition System [6], which was based on the Robocasting process developed by Sandia National Laboratories [7-8].

The second process uses a Micro-Scale Robotic Deposition System [6], which was based on the Robocasting process developed by Sandia National Laboratories [7-8]. This process uses air pressure to extrude material for deposition and uses one-way control valves for starting and stopping of the extrusion process. These “retraction” processes can be applied to the ABEF system by retreating the ram extruder from touching the materialreservoir plunger. This would in theory reduce the applied pressure, thereby stopping material extrusion.In this paper material deposition tests are conducted, where the ram extruder retreating velocity is varied in order to develop a technique that can be used for starting/stopping the extrusion process on demand. A new extrusion mechanism is designed to improve process capabilities.

A new extrusion mechanism is designed to improve process capabilities. Parts are then fabricated with the use of starting/stopping the extrusion on demand with emphasis being placed on deposition of sharp corners and edges.Figure 1: Schematic showing the location of the die land for an extrusion nozzle.Ram AxisZ-AxisMaterial ReservoirY-AxisX-AxisFigure 2: Photograph of the ABEF system used in this study.STOPPING MATERIAL FLOW TESTINGExperiments were conducted to test the feasibility of stopping material deposition on demand by retreating the ram extruder. The schematic of the original extrusion mechanism utilizing a plastic syringe and hypodermic needles is shown in Figure 3. Single lines were deposited at table velocities of 25 mm/s and 30 mm/s with a desired ram force of 308 N at a standoff distance of 150 µm. These process parameters yielded good deposition as determined from previous part fabrication using the original extrusion mechanism.

These process parameters yielded good deposition as determined from previous part fabrication using the original extrusion mechanism. Once the table velocity has reached a steady state and traveled a fixed distance (5 mm), the ram was retreated at a given velocity ranging 0-10 mm/s. After the ram is retreated the amount of excess material deposited was measured from the desired stopping location. Figure 4 shows the deposition length for table velocities of 25 mm/s and 30 mm/s. By examining Figure 4 it can be seen that the extrusion is not able to stop on demand, regardless of the ram retreating velocity. This is due to the fact that the ram axis is not directly connected to the plunger (see Figure 3). When the ram retreats it does not directly stop the pressure being applied to the fluid within the plastic syringe. It only detaches itself from the plunger.

When the ram retreats it does not directly stop the pressure being applied to the fluid within the plastic syringe. It only detaches itself from the plunger. In order for the material to stop extruding, equilibrium must be reached between the applied extrusion force and the shear stress created by fluid motion. This occurs when enough material has been released, decreasing the amount of material present within the same material reservoir volume.

This occurs when enough material has been released, decreasing the amount of material present within the same material reservoir volume. From these results the ram retreating method of stopping the extrusion process was not a viable solution with the plastic syringe extrusion mechanism design.Ram AxisPlungerLoad CellMaterial ReservoirRubber SealNozzleFigure 3: Schematic of original extrusion mechanism utilizing a plasticsyringe and hypodermic needles.140)m m12025 mm/s 30 mm/s( h t g n e L n o i t i s o p e D s s e c x E10080604020 026 Ram Retreating Velocity (mm/s)481012Figure 4: Ram retreating test data for the original extrusion mechanism.NEW EXTRUSION MECHANISM DESIGNIn order to improve the process by adding the ability to stop the extrusion on demand, a new extrusion mechanism was designed on the basis that to stop the extrusion it would be necessary to retract not only the ram axis but also the plunger.

To accomplish this, a housing was placed around the load cell that would connect directly to the plunger. With the load cell enclosed as the ram axis is retreated the load cell touches the housing, thereby raising the plunger. A schematic of the new design is shown in Figure 5.To further improve the system it was decided to replace the plastic syringe and hypodermic needles with a two-piece metal reservoir. The metal reservoir reduces variability in the internal pressure that was present in the original extrusion mechanism due to expansion of the plastic syringe.

The metal reservoir reduces variability in the internal pressure that was present in the original extrusion mechanism due to expansion of the plastic syringe. The mechanical design is such that different dies can be used for the desired material deposition height and width.Figure 5: Schematic of new extrusion mechanism utilizing an interlockingsystem made of 304 stainless steel.To determine a suitable design for a new material reservoir the resulting stress onthe reservoir was calculated usingσ=F A=( R πF 2 o R − i)2where σ is the axial stress acting on the system, F is the extrusion force, A is the equivalent area, and Ro and Ri are the outer and inner radii of the reservoir, respectively. Substituting in the values of F = 4.4 kN (maximum extrusion force), Ro = 15.2 mm, and Ri = 12.7 mm, the maximum axial stress is determined to be σ = 63 MPa.

Substituting in the values of F = 4.4 kN (maximum extrusion force), Ro = 15.2 mm, and Ri = 12.7 mm, the maximum axial stress is determined to be σ = 63 MPa. In order to guarantee the reservoir would not fail under this stress the wall thickness was increased to 12.5 mm and 304 stainless steel was chosen for its yield stress of 500 MPa. With these changes made, a factor of safety of 7.9 is present in the design. Figure 6 shows the newly designed material reservoir as implemented into the ABEF system. For the design of the nozzle, a diameter of 0.5 mm was chosen as a similar diameter (580 µm) to that previously used with the plastic syringe and a die land length of 5 mm was chosen, which was a 50% reduction in length from the previously used hypodermic needles. Figure 7 shows a close-up view of the die land of the new reservoir.

Figure 7 shows a close-up view of the die land of the new reservoir. Tests proved that the new design worked without failure.Load Cell HousingConnectorPlungerMaterial ReservoirFigure 6: 304 stainless steel extrusion mechanism design implemented in the ABEF system.(1)Die LandFigure 7: Picture of new extrusion mechanism die land.The ram retreating experiments were repeated using the new extrusion mechanism. Three different extrusion forces were used: 308, 445, and 582 N. These extrusion forces were chosen due to good deposition results during part fabrication using the original extrusion mechanism. Figure 8 shows the results from the deposition tests using the new extrusion mechanism. The extrusion was able to be stopped on demand using the retreating method at all three extrusion forces.

Figure 8 shows the results from the deposition tests using the new extrusion mechanism. The extrusion was able to be stopped on demand using the retreating method at all three extrusion forces. There was an exponential decrease in the amount of excess deposition as the retreating velocity increased for all extrusion forces.100)m m( h t g n e L n o i t i s o p e D s s e c x E80604020F d F d F d= 308 N= 445 N= 582 N0 00.511.5Ram Retreating Velocity (mm/s)Figure 8: Ram retreating results using the new extrusion mechanism with a table velocity of 25 mm/s and a standoff distance of 0.5 mm.In order to determine the repeatability of stopping the extrusion process, the line deposition tests were conducted 10 times for ram retreating velocities of 0, 0.25, 0.5, 1, 1.5, 2, 2.5, and 3 mm/s at an extrusion force of 308 N. Standard deviations of 0.0, 9.2, 1.5, 4.7, 1.0, 0.0, 0.0, and 0.0 mm were calculated for the ram retreating velocities, respectively.

Standard deviations of 0.0, 9.2, 1.5, 4.7, 1.0, 0.0, 0.0, and 0.0 mm were calculated for the ram retreating velocities, respectively. Figures 9-12 show pictures of the line deposition tests at ram retreating velocities of 0.5, 1, 1.5, and 2 mm/s, respectively. Velocities higher than 2 mm/s were also able to stop the extrusion on demand. From the figures it can be seen that the process is repeatable, with small variations in excess deposition between individual runs at a given ram retreating velocity. Once the ram begins to retreat, the deposited lines decrease in thickness until gaps become present in the deposition. The gapped deposition is due to the reduction in extrusion force. As the extrusion force decreases the ceramicpaste extrudes slower and slower eventually not being able to keep up with the table velocity causing the material to shear thereby causing gaps in the deposition.

Figure 13 shows the results from the repeatability tests.10 mmFigure 9: Three runs of line depositions with ram retreating velocity of 0.5 mm/s and Fd = 308 N. Black line indicates desired stopping location.10 mmFigure 10: Three runs of line depositions with ram retreating velocity of 1 mm/s and Fd = 308 N. Black line indicates desired stopping location.10 mmFigure 11: Three runs of line depositions with ram retreating velocity of 1.5 mm/s and Fd = 308 N. Black line indicates desired stopping location.10 mmFigure 12: Three runs of line depositions with ram retreating velocity of 2 mm/s and Fd = 308 N. Black line indicates desired stopping location.120)m m100( h t g n e L n o i t i s o p e D s s e c x E80604020000.51.5 Ram Retreating Velocity (mm/s)122.53Figure 13: Chart of retreating test data at a table velocity of 25 mm/s, Fd = 308 N, and a standoff distance of 0.5 mm.

Error bars represent one standard deviation.PART FEATURESWith the ability to stop the extrusion on demand, the next objective was to test this technique for part fabrication. The first part used for testing was a single layer square with an interior rectangular hole. In order to fabricate this part a motion path as shown in Figure 14 was used. First the exterior contour was deposited for a smooth outer surface. Next the interior was deposited in a rastering motion. The black lines in Figure 14 indicate material deposition and the gray lines indicate no deposition. Figure 15 shows the fabricated part. The desired deposition dimensions of the part were 25 mm X 25 mm with a 10 mm X 10 mm interior square hole. Ten separate measurements were taken of each dimension after freeze drying to allow the part to reach brown state.

Ten separate measurements were taken of each dimension after freeze drying to allow the part to reach brown state. Measurements are taken in the brown state due to the deposited material having a very low yield stress before drying, allowing for easy deformation during measurement. The average length parallel to deposition was 24.2 mm and the average length perpendicular to deposition was 24.8 mm with standard deviations of 0.4 and 0.6 mm, respectively.This yields a dimensional error of 3.2% and 0.8% for the parallel and perpendicular dimensions, respectively, after freeze drying. The interior hole was 11.1 mm parallel to deposition and 11.5 mm perpendicular to deposition with standard deviations of 0.24 and 0.3 mm, respectively. The dimensional error for the internal holes is 11.1% and 15% for the parallel and perpendicular deposition directions, respectively.

The dimensional error for the internal holes is 11.1% and 15% for the parallel and perpendicular deposition directions, respectively. The error present between the desired and actual dimensions is due to the shrinkage of the part during the freeze drying stage. The differences in size between the parallel and perpendicular measurements are caused by different drying rates of the part. When the part is drying, the parallel deposition direction is allowed to dry more uniformly with all of the deposition ends open to the environment as opposed to the perpendicular direction having only a single line open to the environment for drying. The larger amount of error present in the internal measurement was due to shrinkage in two directions as compared to the single direction with the external dimensions.Figure 14: 2D test part motion path.

Dashed lines indicate no deposition.Figure 15: Top view of deposited material from motion path in Figure 14 at a table velocity of 15 mm/s, standoff distance of 0.5 mm, and Fd = 308 N.With the success of being able to fabricate 2-dimenisonal shapes by starting and stopping the extrusion on demand, the described technique was implemented for the fabrication of 3-dimensional parts. The part fabricated in this test was a slanted block with an interior hole. The same motion path is followed as for the 2D part, but between layers a 200 µm shift is made in both the X and Y directions. Figures 16 and 17 show the side and top view, respectively. As can be seen from Figure 16 some curling is present at the edges of the part. This is caused by the uneven drying rate of the part in roomtemperature. The first couple of layers dry quickly compared to the remaining layers.

This is caused by the uneven drying rate of the part in roomtemperature. The first couple of layers dry quickly compared to the remaining layers. The internal layers are enclosed from the external environment, allowing for more uniform drying.Figure 16: Side view of slanted part at a table velocity of 15 mm/s, standoff distance of 0.5 mm, and Fd = 308 N.Figure 17: Top view of part in Figure 16.SUMMARY AND CONCLUSIONSA new extrusion mechanism for the process of aqueous-based extrusion fabrication of high solids loading ceramics on demand has been designed and implemented, providing an extrusion capability that cannot be done with the use of a plastic syringe for extrusion. Deposition tests were conducted to examine the relationship between ram retreating velocity and excess material deposition. There is an exponential decrease of excess material deposition with increasing ram retreating velocity.

There is an exponential decrease of excess material deposition with increasing ram retreating velocity. The implementation of the new extrusion mechanism allows for the extrusion to start and stop as programmed. The ability to start and stop the extrusion on demand allows for more complex 3D part fabrication including sharp corners and features. Parts were fabricated using the new material reservoir to demonstrate the developed capability. Several features were measured for dimensional accuracy compared to desired dimensions. More shrinkage was present in the parallel deposition areas than the perpendicular deposition areas due to a more even area present for drying. The exterior part dimensions were more accurate than interior due to less surface area for drying.ACKNOWLEDGEMENTThis work was supported by the Air Force Research Laboratory under ContractFA8650-04-C-5704.REFERENCES 1.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAREAL SURFACE CHARACTERIZATION OF LASER SINTERED PARTS FOR VARIOUS PROCESS PARAMETERSP.Delfs* and H.-J. Schmid†Direct Manufacturing Research Center, Paderborn University, Paderborn, Germany † Particle Technology Group, Paderborn University, Paderborn, GermanyAbstractLaser sintered polymer parts consist of rough surfaces due to the layered manufacturing and adherence of incomplete molten particles. The absolute roughness depend on various process parameters like build angle, spatial position, build temperature, exposure order and layer time. Analyses with the help of several areal roughness values of DIN EN ISO 25178-2 considering these parameters are introduced in this paper.

Analyses with the help of several areal roughness values of DIN EN ISO 25178-2 considering these parameters are introduced in this paper. Multiple build jobs with 120 µm layer thickness and PA2200 powder were built on an EOS P396 machine using the same build job design with varying process parameters. An individual sample part was designed to receive lots of surface topography information with optical 3D measurements. The results show roughness dependencies for 0° to 180° build angles in 15° steps and eleven distributed in-plane and three axial direction positions depending on different build temperatures, reversed exposure order and layer times. Limitations of the varied parameters are finally derived for the manufacturing of improved surface qualities.IntroductionIt is well known that polymer laser sintered parts possess rough surfaces. Nevertheless, only a few investigation are focused on objective surface roughness values for additive manufactured parts.

Nevertheless, only a few investigation are focused on objective surface roughness values for additive manufactured parts. Mostly, surface roughness is analyzed with the example of the profile arithmetic mean value 𝑅𝑎. Reeves and Cobb [RC97] introduced surface roughness analyses for stereolithography parts with their widespread “truncheon” sample part which covers build angles from 2° to 180° in 2° steps. Later, Upcraft and Fletcher [UF03] demonstrated in their review article advantages and challenges of several additive manufacturing processes including a surface roughness comparison depending on layer thickness and a build angle between 10° to 90°. Bacchewar et al. [BSP07] introduced an ANOVA study and analysis of process parameter influences on surface roughness. They investigated how the parameters build orientation, laser power, layer thickness, beam speed and hatch spacing influence the roughness with the example of the profile arithmetic mean value 𝑅𝑎.

They investigated how the parameters build orientation, laser power, layer thickness, beam speed and hatch spacing influence the roughness with the example of the profile arithmetic mean value 𝑅𝑎. Build orientation was found to be the significant parameter affecting surface roughness followed by layer thickness for upward directed surfaces and laser power for downward directed surfaces. Correlating results from the metal technology came from Strano et al. [HSE+13] who analyzed and set up a prediction model for laser melting part surfaces of 316L material. They used as well a surface profilometer and a scanning electron microscope for surface analysis. They found in detail an increasing density of spare particles along step edges as the build angle increases.Furthermore, only a handful examination are public where areal roughness values are applied to laser sintered parts. Thompson et al.

Thompson et al. [TSG+17] compared different three-dimensional measurement methods, namely confocal microscopy, coherence scanning interferometry, focus variation microscopy and X-ray computed tomography. They applied measurements to selective laser melted Ti6Al4V part surfaces and analyzed them with various areal texture parameters of ISO 25178-2. Grimm at al. [Gri16][GWW17] used confocal microscopy for an evaluation of the feasibility of three-dimensional surface parameters for quality control in additive manufacturing. Furthermore, they automated the surface analysis via an industrial six-axis robot and applied measurements to laser sintering and laser melting part surfaces.2624This work is focused on the application of areal roughness values to laser sintered part surfaces. A systematic approach with high statistic results was aimed at.

A systematic approach with high statistic results was aimed at. The presented results of this work contain process parameter variations for a fixed machine type, layer thickness, powder batch and exposure parameters and its influences on surface roughness values.Basics and measurement conditionsThe analyses of all the surfaces were done via defined texture values of DIN EN ISO 25178-2. In detail, the values areal arithmetic mean 𝑆𝑎 and areal roughness depth 𝑆𝑧 were considered, where 𝑆𝑎 is defined as follows:𝑆𝑎 =1 𝐴∬|𝑧(𝑥, 𝑦)|𝑑𝑥 𝑑𝑦𝐴𝑆𝑧 is defined as the sum of the highest peak and the deepest valley point of the considered area.The measurements were produced with a “3D-Macroscope VR-3100” of Keyence Corporation. This system operates with fringe light projection as shown in figure 1. A tilted light source projects fringe light pattern onto the sample and a CMOS detector collects the reflected light. Via triangulation, the height information is then derived.

A tilted light source projects fringe light pattern onto the sample and a CMOS detector collects the reflected light. Via triangulation, the height information is then derived. A resulting exemplary surface measurement of a 15° tilted surface is also shown in figure 1, where the stair-stepping effect and adhered particles are obvious.Figure 1: Exemplary 3D-measurement of 15° tilted LS surface (left) and measuring principle of fringe light projection (right).The measurement area of all measurements are approx. 7 x 9 mm². The corresponding 𝑆𝑎 value is calculated by the whole data points. The 𝑆𝑧 value is calculated for nine different sub-areas of 2 x 2 mm² size for each measurement to reduce the influence of individual particle agglomerations. The build angle range of 0° to 180° in 15° steps is covered by the hexagon sample part demonstrated in figure 2. The definition of surface angle directions are also given in the two sketches.

The build angle range of 0° to 180° in 15° steps is covered by the hexagon sample part demonstrated in figure 2. The definition of surface angle directions are also given in the two sketches. 0° is a surface parallel to the building platform and upward directed, 90° is a surface orthogonal to the building platform and 180° is a surface downward directed into the powder bed. The angles are shown with front views of the two necessary orientations the sample part has to build with. The basis of the part is formed by a hexagon with feed sizes of 18 mm and 20 mm, which is extruded for 18 mm. The part is hollowed and consists of a wall thickness of 3 mm. Lastly, the center has a 10 mm sized quadratic extruded cut for reproducible post-processing purposes. The design of the sample part enables the part to stand on every surface while the opposing surface is aligned parallel to the ground.

The design of the sample part enables the part to stand on every surface while the opposing surface is aligned parallel to the ground. This feature allows a constant plane measurement of each surface from the part with the measuring system without the need of a special bracket.2625Figure 2: Two build orientations of hexagon sample part with appropriate build angles.Experimental designFor the manufacturing of sample parts one build job design was prepared, see left image of figure 3. For the analyses covered here, only the blue hexagon parts of the design are of relevance. The whole build space of a EOS P396 machine (350 x 350 x 620 mm³) minus small part free margins on each side is used with sample parts. The hexagon parts are manufactured at three z-height levels, each level with two layers of parts in above- mentioned two orientations. Each layer of parts consists of 22 parts positioned in a 5 x 9 position matrix as shown on the right side in figure 3.

Each layer of parts consists of 22 parts positioned in a 5 x 9 position matrix as shown on the right side in figure 3. The red marked positions are the ones that were measured and analyzed. The dashed lines in red and black indicate center and outside areas. The presented build job was manufactured with various process parameters. Following parameters were held constant for each job:Machine: EOS P395 - Material: PA 2200 mixed powder of one batch - PartPropertyProfile Balance: 120 µm layer thickness - Preheating time: 4 hours - Unpacking temperature: below 50 °CThe results of four different build jobs are included here. The variation contains firstly the build temperature, changed by 2 K, and further a reversed exposure order of the parts and an additional 30 s layer time, which is summarized in table 1. A reversed exposure order changes the duration the molten bath is covered with powder material. In build job #4 a dummy part extends this time by 30 seconds, which is approx.

A reversed exposure order changes the duration the molten bath is covered with powder material. In build job #4 a dummy part extends this time by 30 seconds, which is approx. four times the exposure duration of the hexagon parts.Table 1: Overview of build jobs and their varied process parameters.Build job numberBuild temperature / °CDetailed variation#1172.5#2174.5#3174.5reversed exposure order#4174.5plus 30 s layer time2626Figure 3: Build job design with top view of x-y-distribution of hexagon samples and division in center (black dashed line) and outsidepositions (red dashed line).Results and discussionThe first results show the roughness values 𝑆𝑎 and 𝑆𝑧 in dependency of the build angle and the four different build jobs in figure 4 and 5. Each data point is the mean value of 33 samples for the distinct build job, averaging all spatial positions.

Each data point is the mean value of 33 samples for the distinct build job, averaging all spatial positions. Beginning with the results of 𝑆𝑎 values the flat surfaces (0° and 180°) show the lowest values followed by 90° surfaces. Upward directed surfaces (15° to 75°) exhibit a decreasing roughness with increasing angle correlated to a cosine relationship due to the stair-stepping effect of the layer wise manufacturing principle. This effect is strongly diminished for downward directed surfaces (105° to 165°) where all angles appear to be comparable rough. Explanatory approaches are a small curling effect on each layers edge and the filleting effect due to the laser penetration geometry [BSP07]. The standard deviations of the most points are overlapping, nevertheless a light tendency can be seen for build job #4 comprising the lowest the roughness values of all build jobs. The areal roughness depth do not confirm this trend, nearly all values for each build angle are on the same range.

The areal roughness depth do not confirm this trend, nearly all values for each build angle are on the same range. Rather noticeable are the even more comparable values for the downward directed surfaces and that they are all lower than the roughness depth of 90° surfaces.2627Figure 4: Areal arithmetic mean roughness in dependency of the build angle for all build jobs.Figure 5: Areal roughness depth in dependency of the build angle for all build jobs.2628For a view on spatial dependencies, analyses according the z-height and x-y-distribution are discussed in the following. In figure 6, the z-height dependency is demonstrated exemplarily for build job #3 and the build angles 0°, 15° and 90° as they represent the extreme points in roughness. Per plane the results of eleven samples were averaged, distinct deviations on individual positions were not found.

Per plane the results of eleven samples were averaged, distinct deviations on individual positions were not found. Also, the comparison of different build heights show no differences beyond standard deviations.Figure 6: Areal arithmetic mean roughness in dependency of the Z-height for build job #3.In figure 7, the X-Y-Distribution is shown by separating sample of the center and outside build area as defined in figure 3. Again, the extreme points with 0°, 15° and 90° build angles are selected to exhibit the qualitative tendencies. Measurements of all build heights are averaged as this has no influence like shown before. Distinct differences can hardly be highlighted. Once more, disparities for the in-plane position can not be found while they appear between the different build jobs. Build job #4 possess the smoothest surfaces, which confirms the influence of increased layer time. For 15° build angle also the increased build temperature has a small positive effect on the roughness.

For 15° build angle also the increased build temperature has a small positive effect on the roughness. The small differences between build job #2 and #3 demonstrate the degree of reproducibility.Figure 7: Areal arithmetic mean roughness in dependency of the X-Y-Position for all buildjobs and 0°, 15° and 90° build angle.2629Figure 8: Zoomed 3D-images of 0° surfaces of build job #2 (left) and build job #4 (right) with increased layer time.An approach for the explanation of smoother surfaces for the build job with increased layer time is demonstrated with exemplary 3D-measurements of 0° build angle surfaces of build job #2 and #4 in figure 8. For this flat surfaces only adhered particles and no stair-stepping effect is responsible for the surface roughness. Both shown images have the same color scale and the differences in local topography gradients are obvious. In the left image the particles protrude much more than in the right image.

Both shown images have the same color scale and the differences in local topography gradients are obvious. In the left image the particles protrude much more than in the right image. The increased layer time for the right part surface provides the adhered particles time to melt more into the surface. The adherence of unmolten particles proceed due to the unavoidable contact of the molten bath and particles of the next powder layer. The heat conduction of the material is very low as well as the viscosity just a few degrees over melting temperature. Hence, the degree of particle melt is mostly driven by the available time with IR-heating from above before the next colder layer of powder is deposited and supposed to abort the adherence process.

Hence, the degree of particle melt is mostly driven by the available time with IR-heating from above before the next colder layer of powder is deposited and supposed to abort the adherence process. Figure 9 sketches the progress of particle adherence to a molten bath and a time correlation to the layer times of build jobs #2 and #4.Figure 9: Sketch of particle adherence progress over time after the last laser exposed layer.Conclusion and outlookIn this paper we demonstrated surface roughness analyses of polymer laser sintered part surfaces via areal surface texture values of DIN EN ISO 25178-2. Build orientation, spatial position, build temperature, exposure order and layer time are varied parameters of the investigations. It was shown that the build orientation has the highest influence on surface roughness followed by the layer where longer durations lead to smoother surfaces.

It was shown that the build orientation has the highest influence on surface roughness followed by the layer where longer durations lead to smoother surfaces. Spatial position, build temperature and exposure order play a minor role.Further analyses are in preparation for an extended variation of parameters e g. machine type, wall thickness, part distance, layer thickness and powder quality. These results will give an even more detailed view of influence factors for the surface quality in polymer laser sintering.2630AcknowledgementThe authors want to thank all industry partners of the DMRC as well as the federal state of North Rhine- Westphalia and the University of Paderborn for the financial and operational support within the project “STEP: Surface Topography Analysis and Enhancement of Laser Sintered Parts”.Literature[RC97]Reeves, E. R; Cobb, R. C.: “Reducing the surface deviation of stereolithography using in- process techniques”. Rapid Prototyping Journal, Vol.3, pp.

 Engineering: Pervasive RP activities at Arizona State UniversityAnshuman Razdan, Mark R. Henderson and Dan Collins MC5106, PRISM, Arizona State University, Tempe, AZ-85287-5106 prism@asu.eduAbstract We present a sampling of widely diverse Rapid Prototyping activities at ASU. Through the interdisciplinary PRISM project, RP is firmly involved in three areas: education, research and outreach. Two courses have been developed, one teaches RP over the web and the other Visualization and RP as applied technologies aimed at interdisciplinary graduates and seniors. Researchers and students from Engineering, Architecture and Industrial Design, Fine Arts and Liberal Arts and Sciences and Business actively use RP.

Researchers and students from Engineering, Architecture and Industrial Design, Fine Arts and Liberal Arts and Sciences and Business actively use RP. A recent formation ofthe PARfC (PRISM Advanced Rapid-fabrication Consortium) Consortium with local industry will create a local and formal center for research, education, service and training in Rapid Fabrication.IntroductionDescription ofPRISMPRISM stands for Partnership for Research In Stereo Modeling. PRISM was establishedover two years ago at ASU to promote interdisciplinary research in the areas of 3D Data acquisition, Visualization and Modeling, and Form realization. PRISM researchers come from diverse backgrounds such as Biology, Fine Arts, Archaeology, Anthropology, Computer Science, Mechanical and Industrial Engineering to name a few.

PRISM researchers come from diverse backgrounds such as Biology, Fine Arts, Archaeology, Anthropology, Computer Science, Mechanical and Industrial Engineering to name a few. The common interest among various disciplines in 3D and higher dimensional data sets is the key motivation for PRISM.Description ofPAR.tCMany rounds of discussions and need for a joint RapidPrototyping.and Fabricationresearch between local companies such as Motorola, Allied Signal, Boeing, PADT, Raytheon and Toy The goals of the PARtC consortium are three fold, Le. Education, Research and Outreach.Inc., led to the formation of the PRISM Advanced Rapid-fabrication Consortium.Equipment and Lab setupPRISM currently operates one lab in the Goldwater Center and is in the process of establishing two mores labs including the Rapid Fabrication Consortium facility.

The lab is equipped with high end SGI and NT workstations, Virtual Reality setup,2 Cyberware laser digitizers and three rapid prototyping machines including the.Genisys (plastic), FDM1650 (ABS lab is equipped with commercial software such as and wax) and a lP5 (manual feed paper). I-DEAS, PRO-E, Maya (Alias) etc. and also inhouse software.Withthe formation ofthe consortium, we are looking to add a DTM machine to our lab.We describe in this paper various interdisciplinary activities at PRISM and Arizona StateUniversity concerning Outreach.are categorized below under Education, Research and221EducationThe education component includes providing opportunities for all students from college freshmen through life.-Ionglearners. It means learning both on and off campus, with university and industrial teaching andmentoring. For education in Rapid Fabrication (RF) to be effective it needs both theory and practice. The capabilities must be cutting edge in pedagogy and technology.

For education in Rapid Fabrication (RF) to be effective it needs both theory and practice. The capabilities must be cutting edge in pedagogy and technology. The result will be graduates and practitioners who can bring the most recent ideas and techniques to industry and continue to push the RF envelope. We have several RF education programs already in place. They are as follows:CoursesThree courses were offered specifically in rapid prototyping during the 1997-98 academic year.IIIEE591C - Industrial engineering course combining one-third RP and two-thirds mechatronics. A fullversion of the Rapid Prototyping course will be offered on campus in the fall 1998 as IEE591A.IEE591 - Web based version of the RP course offered to students on the web at Oklahoma State University, Allied-Signal and Boeing Helicopter.ART591 - VIZProto, Interdisciplinary 3D visualization and rapid prototyping course offered to students across disciplines.

Students applied the technology in 3D scanning, modeling and rapid prototyping. Student projects are available at http://surdas.eas.asu.edu/~razdan/Class/3dvizrp.We give three examples from the student projects. The first project was to design and make a functional prototype for a tape measure. The model was designed on Form-Z CAD package and various design parameters were noted. These were compared against the dimensions measured from the rapid prototyped model (see Table 1). The model was built on a Genisys. The differences were attributed to resolution limitation of the RP machine. After only minor finishing the tape measure was functional. Figure 1 is the concept design and Figure 2 is a picture of the actual model. The comparison resulted in characterization of the Genisys machine.222Figure 1: LonClent ••"';"........measureThe second project involved making a scaled prototype.

The comparison resulted in characterization of the Genisys machine.222Figure 1: LonClent ••"';"........measureThe second project involved making a scaled prototype. This project combined the hightechnology of 3D scanning with a more low-tech approach to sculptural prototyping. Initially, an extracted wisdom tooth was scanned using a Cyberware 3D Digitizer. Although much of the tooth scanned with surprising.clarity of detail for an object so small, certain problem areas lacked accurate data. The most notable of these areas was quite close together, creating a visual overlap from the digitizer. Using QuickSlice, the proprietary software from Stratasys, the three-dimensional computer model was first repaired, and then sliced at even intervals. Each of these 40 layers was projected onto styrofoam2 of the rootsthewere traced by223hand. The layers were hand-cut with a hot wire and assembled to form an enlarged replica of the original object. See Figure 3.

The layers were hand-cut with a hot wire and assembled to form an enlarged replica of the original object. See Figure 3. Not clearly visible, but the original tooth is in the glass case on the top left of the image.Design ProjectsThe lab was usedbystudents in both the courses listed above and other course projects offeredthroughout the year. A large number of student design projects were built in the RP lab duringth~·spring1998Sem,ester.••• The.table below shows the number of parts and amount of m,at~rialused. It should be keptin mind that even thOllgh the activity is fairly high, this year was the frrstto have these two machines and tl1eir availability was not generally publicized. One notable senior design project was the assembly of the 48 piece articulated landing gear (See Figure 4).

One notable senior design project was the assembly of the 48 piece articulated landing gear (See Figure 4). Table 2 notestheRP activity during the spring 1998 semester.RPMachine Stratasys GenisysStratasys FDM1650CapabilitiesLower resolution, faster, polyethylene parts Hightwresolution, slower build, ABS, wax and support structure rnaterialsNo. of Parts 70-8088Material Used25(*$100) cartridges2 spools each of modeling ($153/spool) & support material ($75/spool)Table 2:RP fr(]lducti(mSpring 98Figure 4: Two views of the 48-Piece Articulated landing gearPrograms Under DevelopmentWe are proposing to develop the following programs as part of the educational offerings. These include Industry Internships, ASU Internships and Seminars/Conferences.

These include Industry Internships, ASU Internships and Seminars/Conferences. These would facilitate exchange of ideas and rich experience for the students and industry alike.224ResearchCurrent ProjectsThe following paragraphs describe ongoing and completed research projects both within and outside of PRISM in the area of Rapid Fabrication. These activities form the basis for research inPARtC.1) Optimizing Part Orientation for Surface Finish, Gautham Kattethota and Mark Henderson.Weare developing a software tool to calculate the expected surface finish quality of all surfaces on a RP model during the build process. SampleRP parts were fabricated with a variety of surface orientations. The surface finishes were characterized using a Sheffield Surface Scanner at Allied Signal (Tempe). Data have shown an unexpected, yet predictable surface finish based on several build factors.

The surface finishes were characterized using a Sheffield Surface Scanner at Allied Signal (Tempe). Data have shown an unexpected, yet predictable surface finish based on several build factors. The software tool will allow the user, through the use of color codes, to compare desired and possible surface finish specifications during interactive part orientation and to select the optimal build orientation. A paper on this is presented at the SFF 98 conference.2) Reverse Engineering From Scanned Data, Ben Steinberg, Anshuman Razdan, Gerald Farin, NSFThe goal is to reduce and convert the data from a digital laser scan into a NURBS surface for reverse engineering and to create a parametric solid model.

A paper on this is being presented at the SFF 98 conference.3) Scientific Visualization Using Tactile Feedback, Anshuman Razdan, NSF.This is the second year of the two-year NSF funded project to make RP models of microscopic and mathematical images to help visually impaired visualize scientific data. Work continues on transforming various 3D data to RP and incorporating surface texture.Figure 5: An enneper surface and CDROM microscope sampleFigure 5 shows two samples from the project (http://surdas.eas.asu.edu/~tactile).2254) Virtual Sculpting (HERA Project), Carl Dahl, Alyn Rockwood, Geoff Duke, Jeff Dorman, Mark Henderson.Int~ractive shape modification through 3D input to a virtual lump of modeling material creates geometric solids, which can be viewed and prototyped. The user can the shape through hand gestures and produce a plastic prototype or solid model as output.

The user can the shape through hand gestures and produce a plastic prototype or solid model as output. project is aimed at conceptual design (Figure 6).Figure 6: A shape of a horse using Virtual Sculpting5) STL File Repair, Surya Suluh, Mark Henderson and Anshuman Razdan.Using algorithms for redundant vertices, missing and overlapping facets, flipped surface normals and unmatched facet edges, a program was developed to repair damaged or problematic STL files automatically and to review both the errors and results graphically. This was a M.S.E. project (finished May 1998).6) Process Planning For Rapid Machining, Norma Hubele, Chell Roberts, Mark lI-IIo·nrt"....C'rvn NSFThe second 3-year phase of this NSF funded research will develop a tool to provide manufacturing advice to a designer in order to achieve minimum NC machining minimum cost or maximum part quality.

NC machining is another way to do rapid fabrication.7) Metal Spray Deposition for Rapid Fabrication, Ampere Tseng, Various agenciesA new, patentable process is being developed to create quick metallic prototypes using metal spray.8) Preserving the past in 3DA recent collaboration has brought the Archaeological Researchlnstitute (ARI) partnership with PRISM to create three-dimensional database ofarcheological ARI at ASU has been the keeper of Native American artifacts thatwere excavated during the Central Arizona Project. Recently, Congress passed the Repatriation Act, which entitles the Native American tribes to gain possession of these artifacts and rebury them. This will result in the permanent loss of a great amount of scientific data whose only record would be through photographs and subjective description.

This will result in the permanent loss of a great amount of scientific data whose only record would be through photographs and subjective description. However, with the joint effort artifacts will be digitally scanned before returning to tribes thus preserving true We have successfully reproduced scaled model replicas using the (see Figure 7).data set. equipment at PRISM2269) The Capitol Project.The original idea to create a model of the State Capitol Building carne out of recentFigure 7: Archaeological artifacts reproducedFigure 8: The Capitol Projectinvestigations into the Capitol Mall district and possible urban renewal projects. The Capitol Building model was built to symbolize a renewed commitment and to celebrate its 100 year anniversary. Architectural plans from 1898 were loaned by the Arizona Department of Administration. Details that could not be translated from the plans were digitally photographed or measured directly off the building.

Details that could not be translated from the plans were digitally photographed or measured directly off the building. The computer model was created using FormZ solid modeling software. Several iterations of the model were produced on the two rapid prototyping machines. A scale model of the original Winged Victory figure that adorns the top of the Capitol was made available by the Arizona State Archives and was scanned in and prototyped. The two models were then translated into bronze casting via silica shell casting at the ASU Art foundry.The Research RoadmapThe following high impact areas have been recognized for emphasis on RF research: Metal fabrication, Non-metal Fabrication, Hybrid Materials, Geometry issues and Prototype Size/Scale.

The following table (Table 4) includes an attempt to categorize the ongoing projects enumerated in section Current Projects.RankResearch Topic (Top 5)Specific Issues for Each TopicOngoing Pro.iects1Metal Fabrication and Rapid ToolingPrecision tooling; direct injection molding tools; dimensional stability and accuracy, material uniformity and properties7,82Issues of ScaleAbility to produce nano- through mega-size parts; minimum feature1,3,4227345Non-metal part materials including ceramicsHybrid and composite materialsGeometryICAD and file format issuessizes; part joining; removing need for segmentation Predictable and tunable material properties; Increased range of materials; Dimensional stability and accuracy Fabricating non-homogenous materials; intentionalinclusion of varied material features; combining fabrication and assembly Self-repairing STL files; Pursuit of STL alternative format; Optimal discretization; ••Software for optimal fabrication parameters including orientation3,4,6,7,81,2,4,5Table 1: The High Priority Research IssuesOutreachOutreach is the process in which ASU shares knowledge and technology transfer with theconsortium members and other partners.

We envision the consortium to have an effective outreach program with the complete supply chain including Large Corporations, Small and Medium Sized Companies, Company Vendors, Rapid Fabrication Equipment Manufacturers and the university. The service is centered around both Large Corporate Members (LCM) who have already invested and adopted RP and Small and Medium size Enterprises (SME) who want to become involved. Service Bureaus are also important in PARtC. Many SMEs are interested in the technology but are turned off by the investment, both financial and staff costs.Many industries in the region are involved in manufacturing and other activities thatcould benefit from incorporating RF. Some are not yet aware of the capabilities or the basics of the technology. For example, Toy Time Toys, Inc., a small entertainment concept design company recently interacted with PARtC to create the next generation of a sporting item to be unveiled during SuperBowl 1999.

For example, Toy Time Toys, Inc., a small entertainment concept design company recently interacted with PARtC to create the next generation of a sporting item to be unveiled during SuperBowl 1999. The design and prototype were taken on as a student project. This was the company's first experience with 3D CAD design and RP. The company is convinced of the advantages the technology provides.AcknowledgementsWe wish to acknowledge the supporters of PRISM and PARtC. In particular, John Macintosh, Mary Bates, Mary Marzke, Ben Steinberg, Gerald Farin, the HERA team, Gautham Kattethota, Jacob, Shay and ARI. Thanks to Allied Signal (Don Benjamin, Doug Deye and Don Deptowicz) and other corporate members.228.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceARTIFICIAL INTELLIGENCE-ENHANCED MULTI-MATERIAL FORMMEASUREMENT FOR ADDITIVE MATERIALSP. Stavroulakis*, O. Davies*, G. Tzimiropoulos†, and R. K. Leach*Manufacturing Metrology Team, University of Nottingham, Nottingham, UK †School of Computer Science, University of Nottingham, UKAbstractThe range of materials used in additive manufacturing (AM) is ever growing nowadays. This puts pressure on post-process optical non-contact form measurement systems as different system architectures work most effectively with different types of materials and surface finishes.

This puts pressure on post-process optical non-contact form measurement systems as different system architectures work most effectively with different types of materials and surface finishes. In this work, a data-driven artificial intelligence (AI) approach is used to recognise the material of a measured object and to fuse the measurements taken from three optical form measurement techniques to improve system performance compared to using each technique individually. More specifically, we present a form measurement system which uses AI and machine vision to enable the efficient combination of fringe projection, photogrammetry and deflectometry.

More specifically, we present a form measurement system which uses AI and machine vision to enable the efficient combination of fringe projection, photogrammetry and deflectometry. The system has a target maximum permissible error of 50 μm and the prototype demonstrates the ability to measure complex geometries of AM objects, with a maximum size of (10 × 10 × 10) cm, with minimal user input.Keywords: Artificial intelligence; data fusion; photogrammetry; fringe projection; deflectometry; segmentation network;IntroductionThe aerospace, automotive and medical industries are realising the benefits of using non- contact optical three-dimensional (3D) measurement over traditional stylus-based contact methods to control their processes and perform dimensional quality control on their products [1],[2].

Depending on the technology used, non-contact optical measurement can include most if not all of the following benefits: it is non-destructive, has improved speed, increased area of acquisition per measurement cycle and reduced cost [3]. For non-contact optical camera-based techniques especially, the ability to leverage the benefits of readily available AI-based image recognition algorithms to enhance the measurement even further provides an additional benefit over stylus- based contact techniques [4]. Non-contact optical techniques, however, are not a panacea, and for some measurement scenarios, contact stylus-based techniques still need to be used, as optical techniques suffer from lower accuracy, material sensitivity and much higher requirements for data handling and storage [3].In optical systems, material sensitivity and measurement range can be minimised by changing optical design and hence many different solutions exist [3].

Consequently, there is a large number of different optical techniques for dimensional measurement which have been developed to work at different length scales and for different applications [5]. The selection of the appropriate technique depends on the specific object size, target measurement accuracy and the material of the object that needs to be measured.One way to improve functionality of 3D measurement instruments is to combine multiple techniques in one setup; however, this requires that the data be combined in one frame of reference,187which adds to the processing time and complexity of the system. In this work, we perform data fusion by using a common reference camera to employ three different camera-based measurement techniques, hence data fusion is performed without the need to register the frames of reference to each other. The techniques employed here are fringe projection (FP), deflectometry (DM) and photogrammetry (PG).

The techniques employed here are fringe projection (FP), deflectometry (DM) and photogrammetry (PG). These 3D measurement techniques nominally function at a 100 mm to 1 m operating distance and at a 10 μm to 100 μm accuracy [6]. The techniques selected are ideal for many of the object sizes and material types which need to be measured in automotive, aerospace and additive manufacturing [6].BackgroundThe three chosen measurement techniques operate via different principles and are largelycomplementary to each other. The principles of operation of all three are analysed in this section:Fringe projectionFigure 1. Schema of a fringe projection setupFringe projection (Figure 1) is a structured light technique which uses a projector and a camera to create a 3D point cloud of a scene observed by a camera either by triangulation (projecting digital fringes) or phase-stepping (sinusoidal fringes).

It is used mainly for measuring diffusely reflecting objects [7].DeflectometryFigure 2. Schema of a deflectometry setup188Deflectometry (Figure 2) is structured light technique similar to fringe projection but instead of projecting the fringe pattern, it is displayed on a TFT screen. This configuration works best for reflective objects, as a bundle of rays from each pixel will reflect from the object into the camera via direct reflection. The types of patterns which are usually used are similar to fringe projection [8]PhotogrammetryFigure 3. Schema of a photogrammetry setupPhotogrammetry (Figure 3) is a technique which creates a 3D point cloud by taking advantage of unique surface features that exist in two or more images of the object taken from different points of view [9].

Multiple camera views are used in order to reconstruct the 3D surface of an object by searching for pixel correspondences in the images taken.In this work, the fringe projection and deflectometry systems use phase stepping (project sinusoidal patterns). The photogrammetry system performs triangulation via feature matching [10]. Fringe projection is most effective on featureless objects, whereas photogrammetry is most effective on objects with a large number of features which can easily be detected by computer vision techniques (e.g. SIFT algorithm) in the image. In terms of speed, photogrammetry usually requires a larger number of calculations to perform dense reconstruction and is, therefore, slower in reconstructing the full 3D point cloud than the other two methods. Lastly, deflectometry is most effective for highly reflective objects in comparison to the other two techniques, which require diffusely scattering objects.

Lastly, deflectometry is most effective for highly reflective objects in comparison to the other two techniques, which require diffusely scattering objects. Hence, by combining these three techniques, 3D surface point clouds can be acquired for reflective and diffuse objects which can be featureless or have multiple features on the surface.Data fusionTo combine the data from all three 3D measurement techniques, we used the same reference camera for all three systems (Figure 4) and a spatio-temporal data collection scheme to collect data from different parts of the image at different times. This scheme has two parts: an image segmentation strategy and a temporal synchronisation strategy.189Figure 4. Schema of the combined setup createdImage segmentation strategyThe image segmentation strategy which was selected employs deep learning through a segmentation network (SegNet) and was trained to identify different materials in an image of the scene taken by the reference camera.

Matlab was used as the framework for building, training and testing the network. The network used was an uninitialised semantic image segmentation network based on the SegNet architecture [11], consisting of encoding layers that downsample the input image by a factor of four and corresponding decoding layers that upsample back to the input size; the RGB images had 201×152 pixels to keep the original aspect ratio whilst being small enough to fit the mini-batch into GPU memory during training. Three material classes were trained by acquiring 400 images of each material, and ground truth images of these were obtained by removing the green background through colour thresholding. The class weighting of the final classification layer used inverse frequency weighting of the total pixel count of each class in the ground truth images.Of the 1200 images, 81% were used for training, with 9% set aside for validation and 10% for final testing.

Random image augmentation was also used to artificially increase the number of training images, consisting of random rotations up to 360°, random reflections horizontally and vertically, and random translations by up to ten pixels horizontally and vertically. The network was trained using an initial learning rate of 0.0005 and a mini-batch size of thirty-two images.In this manner, all the pixels in an image could be labelled by predicted material class using the trained network and a mask for each was created (Figure 5a-b). As the reference image pixels have essentially been classified according to material, the point clouds could now be created only in specific regions by implementing the correct image mask for each measurement method.

Therefore, to optimise the measurement success and associate each measurement technique with a specific material, the Ti6Al-4V mask was used when measuring via photogrammetry (Figure 5c), the carbon fibre mask was used when measuring with deflectometry (Figure 5d) and the polymer mask was used when measuring with fringe projection (Figure 5e).190(a)(b)(c)(d)(e)Figure 5. Results of image segmentation of a scene containing polymer, titanium, and carbon fibre objects. Showing (a) initial image, (b) image segmentation overlay results, and three masks created for (c) titanium, (d) carbon fibre, and (e) polymerTemporal synchronisation schemeThe temporal synchronisation scheme used consists of cycling all three measurement sub- systems (fringe projection, deflectometry and photogrammetry) in the time domain by activating the required modules used by each technique (projector, TFT screen, rotation stage) at the appropriate time and taking measurements of the scene.

Before cycling the measurement techniques though, an image of the scene is taken in order to segment the regions of the image which will be used by each technique. Thereby three measurement masks are made and are used by each technique in turn to reconstruct a selected group of pixels which have the best chance of working and ignore the rest.As all three point clouds are taken by the same reference camera, they are automatically placed in the same frame of reference. The partial 3D data can, therefore, be automatically fused together to form an optimised 3D point cloud of the scene containing all three objects. This191temporal synchronisation strategy allows for both optimised data fusion, minimisation of measurement data and reduction of noise, as each measurement technique works best for a specific material.ResultsSome of the challenges encountered during the image segmentation strategy were primarily related to spurious reflections and ambient light changes.

Below we have identified three cases which needed to be taken into account when training and segmenting the images.Case 1: Reflections from objects in the sceneIn Figure 6, the effect of reflections from objects of a different class is shown. Regions of the carbon fibre object are confused for being parts of the metal material class.(a)(b)Figure 6. Mistakes in prediction caused by reflections of other objects from the carbon fibre surface with (a) initial image and (b) image segmentation overlay results.Case 2: Reflections from environmental lightingIn Figure 7, the effect of ambient lamp reflections is shown. Regions of the carbon fibre object are confused for being parts of the polymer and titanium material class.(a)(b)Figure 7.

Regions of the carbon fibre object are confused for being parts of the polymer and titanium material class.(a)(b)Figure 7. Mistakes in prediction caused by ambient lamp reflections from the carbon fibre surface with (a) initial image and (b) image segmentation overlay results.192Case 3: Illuminating objects with different lighting schemeIn Figure 8, the effect of applying a different lighting scheme to that used in the original training scheme is shown. Regions of the carbon fibre object are confused for being a part of the background, and the shadow from titanium and polymer objects are confused for being a part of the titanium material class.(a)(b)Figure 8. Mistakes in prediction caused by illuminating the objects with a different lighting scheme (projector) to that used during the training phase (ambient room lighting).DiscussionIn this work, an investigation into detecting different AM object materials via use of AI segmentation has been presented.

Specifically, the detection of Ti-V6-Al4, nylon polymer and carbon fibre were investigated. Several problems were identified in the segmentation of the images, which were related to the discrepancies in lighting and reflections which existed when acquiring images from the experimental setup.

In particular, reflections from other objects, reflections from ambient light sources and the significant change in lighting create errors in the segmentation of the images by labelling a large proportion of pixels erroneously.Future work will include controlling the lighting and reflections used when creating the images by shielding the samples from ambient lighting as well as training the network by including more illumination conditions in order to improve robustness and finally customizing the selection of network parameters to enable more accurate segmentation of the images.AcknowledgementsThis work is sponsored by EPSRC grants EP/M008983/1 and EP/L016567/1 and the EU Framework Programme for Research and Innovation – Horizon 2020 – Grant Agreement No 721383.References[1]J. E. Muelaner and P. G. Maropoulos, “Large volume metrology technologies for the light controlled factory,” Procedia CIRP, vol. 25, pp. 169–176, 2014.[2] K.

K NOT WHAT ADDITIVE MANUFACTURING CAN DO FOR YOU…I. GibsonEngineering Design and Innovation Centre, National University of Singapore, Singapore 117576 and Centre for Rapid & sustainable Product Development, IP Leiria, PortugalAbstractThe paraphrase of John F Kennedy’s famous words is for 2 purposes. Firstly it is to acknowledge that there are some people who have considered that it is a major part of their life’s work to promote Additive Manufacturing (AM) technology as primarily a selfless act. AM comprises an outstanding range of technology that should be brought to public attention as a true revolution in how we design and manufacture products. The second purpose is to show that technology development is only one part of this promotion process and that there are other ways in which we can get involved.This paper describes the author’s journey over the (approximately) 20 years since he was introduced to what was then called Rapid Prototyping (RP).

It is not a catalogue of research and development projects but rather a list of activities that he has been involved in to help promote and support AM technology over these years. It will describe the conferences, activities, associations and publications that have been created to allow academics and professionals to describe and discuss their work amongst themselves and to the larger society.IntroductionThe motivation for preparing this paper has much to do with the Solid Freeform Fabrication Symposium. Prof. Dave Bourell and his associates at UT Austin are a prime example of the underlying message. If it were not for the selfless organisation of the SFF over all these years, the whole AM community would be considerably poorer. This year (2012), one new objective of this symposium is to explore a new area within the (normally very research oriented) SFF with the assistance of Terry Wohlers, of AM in the wider community.

This year (2012), one new objective of this symposium is to explore a new area within the (normally very research oriented) SFF with the assistance of Terry Wohlers, of AM in the wider community. Terry is another prime example of selfless contribution to the AM community. This is not to say that Terry and Davie are completely unknown and uncelebrated for their contributions, but we should also acknowledge that the AM community is very fortunate to have both of these individuals.It is not the author’s intention to try and place himself amongst these lofty individuals; that is for others (and time) to determine. The intention is that by showing how he got involved in different activities, then anyone can. The approach is to follow a timeline that covers his movement over the last 20 years from Nottingham, where he first learnt about Rapid Prototyping, to Singapore where he is now.

The approach is to follow a timeline that covers his movement over the last 20 years from Nottingham, where he first learnt about Rapid Prototyping, to Singapore where he is now. It is a personal account that pays much tribute to those individuals the author has had the privilege to work with. There have been many more and93no apology is made for omission, but those who are named have had a significant impact on his career in AM.The Early Years: Nottingham (1991-4)Tribute must be specifically paid to Phill Dickens, who introduced the author to Rapid Prototyping and who was highly instrumental in the early development of the technology within the UK manufacturing industry, and to a certain extent within much of Europe as well. The latter in particular (in the early years at least) relates to the ‘Nottingham Conference’ (European Rapid Prototyping Conference), which he set up and was the first regular Rapid Prototyping conference in Europe.

In the early days this conference was like a hybrid of the SFF and the SME Rapid conference, combining the best of both elements and latching on to the fact that industry and application was key to the sector growth in Europe, rather than the technology itself (like perhaps in the US). Whilst there are a few notable technology milestones in Europe, the European industry has really leveraged on US technology.The Nottingham conference (of which the author helped organise the first two and which was not always held in Nottingham), served the purpose of providing a focus for those wishing to research into RP and for industry to prove that there was a future for the technology. Europe therefore developed an outstanding reputation for application-based systems for tooling, casting and short-run production solutions; in some ways even better than those in the US.

Europe therefore developed an outstanding reputation for application-based systems for tooling, casting and short-run production solutions; in some ways even better than those in the US. After a few years, the importance of this conference dwindled: Euromold took over as the main European industry show and there became numerous outlets for academic work. It was not the prevailing influence in academia that the SFF was, but it certainly served its purpose.For academics, journal output has always been a vital performance indicator and with this new manufacturing technology the question started to be asked about where to publish. Certainly there are numerous high quality journals that can provide an outlet for research into materials, processing, software and management.

Certainly there are numerous high quality journals that can provide an outlet for research into materials, processing, software and management. However, when still at Nottingham University, Ian Campbell and the author decided that since there appeared to be an increasing interest in working with these machines, there was definitely room for a journal of Rapid Prototyping.We investigated a few possible publishers and quickly opted for MCB University Press, who at the time were trying to expand their technology range and who were also very keen on having us on board. I particularly liked the Journal of Integrated Manufacturing Systems (edited by David Bennett), which was one of their titles, and modeled much of the Rapid Prototyping Journal on this. The only real impediment to the journal stemmed from the fact that MCB, who later became Emerald Publishing, were not used to science and engineering journals.

The only real impediment to the journal stemmed from the fact that MCB, who later became Emerald Publishing, were not used to science and engineering journals. It therefore took some time to convey on them the importance of an SCI impact factor and it took more than 10 years before they relented. Whilst the impact factor has see-sawed around 1.0 (which is quite high for a manufacturing journal), the RPJ has been very successful and is firmly established as the top journal for Additive Manufacturing. Readers may be interested to note that the H-Index for the journal is presently around 45, which is extremely healthy.

Readers may be interested to note that the H-Index for the journal is presently around 45, which is extremely healthy. Whilst there are a number of journals now focusing on AM (Journal of Rapid Manufacturing, Virtual and Physical Prototyping, etc.), which are all good journals and fill important gaps in the publishing sphere, it is hard to see why the RPJ should not remain the main outlet provided Ian, Dave (the most recent94US editor after Brent Stucker, who joined a few years after the journal was established, to help deal with the increasing paper load, allowing us to assume regional editor roles) and the author maintain a good ship.The Formative Years: Hong Kong (1994-2005)Transferring to Hong Kong University was, like many career choices, a serendipitous rather than calculated move. It turned out that there was little interest in RP at the time of arrival in 1994 but there was excellent support for anyone prepared to put in the effort.

It turned out that there was little interest in RP at the time of arrival in 1994 but there was excellent support for anyone prepared to put in the effort. This was how the author came to acquire one of the first Selective Laser Sintering (SLS) machines in Asia. The author’s enthusiasm for RP was quite intense at the time and the connection to the RPJ enabled him to quickly develop a reputation within Asia.In 1997 the author and his colleagues organised the 2nd Int. Conf. on Manufacturing Automation, ICMA. This allowed them to bring Terry Wohlers over to Hong Kong to present a keynote. A conversation with Terry at the time covered how it may be possible to promote and develop the RP community. The author had just been involved in establishing the Hong Kong Rapid Prototyping and Tooling Association, which aimed to promote the technology in the region.

The author had just been involved in establishing the Hong Kong Rapid Prototyping and Tooling Association, which aimed to promote the technology in the region. Considering that this association would have grown faster and with greater credibility if it were somehow connected to similar organisations around the world, the idea of GARPA was born.The Global Alliance of RP Associations has served a purpose in adding value to conferences and allowing fledgling associations to leverage from overseas connections to make it easier to establish themselves. Completely free, voluntary and non-profit, its operations have been limited but nonetheless useful. Starting with around 11 member associations, GARPA has grown to include around 25 countries.

Completely free, voluntary and non-profit, its operations have been limited but nonetheless useful. Starting with around 11 member associations, GARPA has grown to include around 25 countries. Terry Wohlers has been the main instrument of this development but special mention should be made to Deon de Beer (South Africa), Paulo Bartolo (Portugal), Alain Bernard (France), and Olaf Diegel (New Zealand), who have all devoted significant amounts of time and effort to supporting GARPA over the years.It is probably at this time that the benefits of promoting AM through writing as well as organising activities should be mentioned. Connections with Terry have allowed the author to write about AM in Asia in the Wohlers Report over the years. Increasing interest in this region due to the strengthening economies has made this article more important in recent years. Whilst at Hong Kong University the author organised and wrote 2 research texts that explored specialist RP technology areas.

Whilst at Hong Kong University the author organised and wrote 2 research texts that explored specialist RP technology areas. The first one was on Software Solutions for RP, which at the time was perhaps somewhat overlooked in importance with most people focused on developing materials and processes. Later, medical applications were emerging in importance and so a text on Advanced Manufacturing Technology for Medical Applications was written. Both these texts were developed using a workshop format, where chapter authors gathered together (the software book workshop was in Hong Kong but the medical book was held in Singapore) to discuss how their work would combine with others to form a successful book. The workshop itself can prove to be a rewarding experience, even without the resulting book.95The Maturing Years: Singapore (2005-)With advancing years you tend to feel an obligation to support those causes and people that you feel deserve it and your attention.

One thing that seemed to missing from all the literature available on AM was an effective teaching text on the topic. Whilst there have been some reasonable attempts at producing such texts, there was nothing aimed at high-level graduate and postgraduate education. Brent Stucker, Dave Rosen and the author all found that they were working on a similar teaching text at roughly the same stage of completion. It seemed logical to combine efforts, which resulted in the Additive Manufacturing textbook that seems to have become quite successful with approximately 1000 copies sold and over 65 academic citations to date according to Google Scholar.The author has been involved as a speaker in many AM conferences over the years. Besides the SFF, he has regularly attended the regular biennial European conferences held in Portugal and Slovenia and the yearly South African RAPDASA conference.

Besides the SFF, he has regularly attended the regular biennial European conferences held in Portugal and Slovenia and the yearly South African RAPDASA conference. The people involved in these conferences have the same drive and commitment of Phill Dickens and Dave Bourell to ensure that AM is at the forefront, with high quality presentations and stimulating dialogue.The collaborations with Paulo Bartolo in Portugal and Deon de Beer in South Africa have led to long-term involvements with their respective institutions. The author is proud to be a visiting researcher at the Centre for Rapid and Sustainable Product Development at IP Leiria in Portugal and will shortly become a visiting Prof. at Vaal University of Technology, South Africa. Both these institutions are worthy of support.

at Vaal University of Technology, South Africa. Both these institutions are worthy of support. Despite the relatively low academic ranking of these institutions (according to worldwide published metrics – not always the best way to judge an institution), both are involved in ground-breaking work, particularly the medical and other research at IP Leiria and the community development programmes using low-cost AM technology at the South African I2P Labs, pioneered by Deon de Beer.This concept of AM supported laboratories for the general public is something the author first saw in South Africa more than 5 years ago and has fully supported since. Attempts have been in vain to implement something like this in Singapore for a long time. However, changes in engineering teaching philosophy at NUS has at least achieving part of this goal.

Attempts have been in vain to implement something like this in Singapore for a long time. However, changes in engineering teaching philosophy at NUS has at least achieving part of this goal. Whilst not available to the general public, the Engineering Design and Innovation Centre allows engineering students from any discipline and year to access this technology in a product and systems development programme called the Design Centric Curriculum.ConclusionIn summary, it is possible to get involved in supporting the AM community in a variety ofways: -Journals: whilst it may not often be possible to start something like the RP Journal, it is possible to become a reviewer. Willingness to review leads to guest editorships and invitations to be Editorial Advisory Board members.Books: there are an increasing number of books being written.

Willingness to review leads to guest editorships and invitations to be Editorial Advisory Board members.Books: there are an increasing number of books being written. Adding chapters is a common approach if you do not have enough material for a full book.96Conferences: it is becoming common to not attend conferences and focus on journal articles. I maintain that it is vital to establish a reputation through conference attendance as part of a career development strategy.Centres: AM is a multidisciplinary field that must also involve industry at a close level. Development of resource centres of excellence using AM as a focal technology is an excellent way to promote its versatility.

Development of resource centres of excellence using AM as a focal technology is an excellent way to promote its versatility. Being involved in centres in other countries can also be a mutually beneficial activity.Blogs, Facebook, LinkedIn, etc.: Whatever you call them, you could take advantage of the various methods of mass social communication to reach out to the wider community as a form of AM ambassador.The importance of associating yourself with other active individuals in the field cannot be stressed enough. This is generally done by agreeing to help each other in some way, like writing joint articles, collaborating on projects or just promising to attend each others’ conference or workshop activities.

This is generally done by agreeing to help each other in some way, like writing joint articles, collaborating on projects or just promising to attend each others’ conference or workshop activities. Whilst it does help to have a team at your home institution, one can make significant contributions as an individual.Of course, the critical activity will always involve quality research work, direct collaboration with industry, and (at least for tertiary education institutions) teaching. Being open and willing to get involved in ‘indirect’ activities is also a very important secondary process that may ultimately reap large rewards if you stick at it long enough.References1. I. Campbell, I. Gibson, D. Bourell, Rapid Prototyping Journal, Emerald Publishing 2. Global Alliance of Rapid Prototyping Associations (GARPA), www.garpa.org 3. I. Gibson (ed.), Software Solutions for Rapid Prototyping, Professional Engineering Publishing (PEP) Ltd. (marketed by John Wiley from 2005), Aug.

sembled Additive Manufacturing – A Hybrid Fabrication ProcessInspired by Origami DesignDongping DengYong Chen*Daniel J. Epstein Department of Industrial and Systems Engineering University of Southern California, Los Angeles, CA 90089, U.S.A.Corresponding author, Phone: (213) 740-7829, Fax: (213) 740-1120, Email: yongchen@usc.eduABSTRACTInspired by the recent developments on origami structures, we investigated a fast AM process for fabricating prototype models of thin-shell shapes. By combining the origami design and the additive manufacturing technology, a new fabrication process named Assembled Additive Manufacturing (AAM) is developed. In the process, a digital model is first analyzed to determine geometries that are fabricated by the layer-based and origami-based approaches. The thin shell model for the origami-based approach is then unfolded into a foldable 2D sheet.

The thin shell model for the origami-based approach is then unfolded into a foldable 2D sheet. After the fabrication and folding of the 2D sheet, a postprocessing process is developed before the final prototype is fabricated using the layer-based fabrication process. Algorithms of analyzing and unfolding polygonal model and the details of the process are introduced in the paper; design parameters, mechanics analysis and test cases are also discussed.KEYWORDS: Additive manufacturing, in-situ assembly, thin-shell structure, origami design, folding, fabrication speed.1INTRODUCTIONObjects with thin-shell structures are commonly used in various products (two product examples are shown in Figure 1a and 1b). Due to the good properties such as lightweight and less used materials, thin- shell structures will be widely used in future engineering systems to satisfy the increasingly important sustainability requirements such as reducing material consumptions.

A general method based on offsetting polygonal models was presented to convert a solid object into a hollow one to reduce its weight [1]. A 3D texturing mapping method was also developed for designing complex internal structures to achieve both lightweight and required strength [2]. Two hollow structure examples are shown in Figure 1c and 1d.Figure 1: A hollow structure example [2].Hollow objects with thin-shell structures are usually difficult to be fabricated by the subtractive- or deformation-based manufacturing processes (e.g. machining and injection molding) due to the difficulty for tools to access the internal features. In comparison, the layer-based additive manufacturing (AM) processes can overcome such accessibility limitations by converting a 3D model into a set of 2D layers for its fabrication [3, 4].

In comparison, the layer-based additive manufacturing (AM) processes can overcome such accessibility limitations by converting a 3D model into a set of 2D layers for its fabrication [3, 4]. Hence, the AM processes seem more suitable for the fabrication of hollow objects with thin-shell structures.A large thin-shell feature usually requires a large number of layers in the AM processes. However, the fabrication efficiency of each layer is not high since only a small amount of material is accumulated in each layer for such thin-shell features. In recent years, a paper folding art named origami has attracted increasing attentions. The origami design is to study how a 2D paper can be folded into a 3D structure [5]. Such a folding-based approach can potentially be a simple and fast method to create hollow objects with thin-shell structures.

Such a folding-based approach can potentially be a simple and fast method to create hollow objects with thin-shell structures. Various issues such as the crease design and folding sequence have been studied in the origami design [6-8].Inspired by the recent developments on the origami design, a new AM process named Assembled Additive Manufacturing (AAM) was developed. The key idea of the AAM process is to integrate the layer-based fabrication approach with appropriate folding operations to significantly improve the fabrication speed.

The key idea of the AAM process is to integrate the layer-based fabrication approach with appropriate folding operations to significantly improve the fabrication speed. Hence the AAM process is a hybrid fabrication method by integrating both folding and layer-based fabrication approaches.(a) Designed ObjectZContinue building top layersBuild layer by layerFolding happensBuild bottom layers(b) Layer-based AM(c) AAMFigure 2: An illustration of the AAM process versus the traditional layer-based AM processes.Figure 2 shows an illustration example to compare the AAM and the traditional layer-based AM processes. Suppose a computer-aided design (CAD) model with thin-shell structures is given (refer to Figure 2a). Based on it, the fabricating process based on the layer-based AM approach is shown in Figure 2b. Note that there may be hundreds of layers in building the middle thin shell portion.

Based on it, the fabricating process based on the layer-based AM approach is shown in Figure 2b. Note that there may be hundreds of layers in building the middle thin shell portion. In comparison, the AAM process uses a hybrid fabrication approach, which is shown in Figure 2c. First, an unfolded structure is fabricated by the layer-based fabrication approach; the fabricated 2D structure is then folded into a 3D structure in a single operation; finally, the layer-based approach is used to continue the fabrication of the remaining top portion. Due to the folding operation, the required building layers can be significantly reduced. Consequently, the AAM process can speed up the fabrication of such hollow objects.2 ASSEMBLED ADDITIVE MANUFACTURING PROCESSThe AAM process integrates the folding and the layer-based fabrication process. An unfolded 2D structure includes designed creases and 2D thin-shell structures. Both of them can be fabricated using the layer-based AM processes.

An unfolded 2D structure includes designed creases and 2D thin-shell structures. Both of them can be fabricated using the layer-based AM processes. The folding of the fabricated 2D thin-shell structures can be applied during the AM fabrication process or after completing the AM fabrication process. Both scenarios are demonstrated in the paper. In addition, multiple materials may be required in the AAM process in order for the creases and 2D thin-shell structures to be properly folded.Various AM processes can be used in fabricating 2D origami sheets and other 3D features in the AAM process. In this paper, the AM process that is used for demonstrating the AAM process is the mask-image-projection-based stereolithography (MIP-SL) process [9, 10]. In the process, a mask image is dynamically defined by a digital micromirror device (DMD). A DMD is a micro-electro-mechanical system (MEMS) device that enables one to simultaneously control over 1 million small mirrors.

A DMD is a micro-electro-mechanical system (MEMS) device that enables one to simultaneously control over 1 million small mirrors. Accordingly a pixel can be turned on or off at over 5 KHz. Using the technology, the energy input related to a mask image can be provided in solidifying a layer of photo-curable resin.Figure 3 shows a bottom-up projection based MIP-SL process. In our study, it is used as the experimental setup of the AAM process. In the building process, a mask image is projected at the bottom of the resin tank. A layer of photocurable resin is cured into the shape defined by the related image. Using a two-way movement method [11], the cured layer can be detached from the coated tank to enable the Z-stage to move up a small distance related to a given layer thickness.

Using a two-way movement method [11], the cured layer can be detached from the coated tank to enable the Z-stage to move up a small distance related to a given layer thickness. Since a mask image can simultaneously cure a whole layer of the resin, the MIP-SL process can fabricate a large 2D origami sheet defined in the AAM process quickly.Z stageComputerPlatformResinCoatingTankCured layersDMDMirrorMirrorimageFigure 3: An illustration of the MIP-SL process.In the AAM process, the foldable portions of a given 3D model need to be identified first. Based on them, appropriate 2D origami structures will be designed. In addition, appropriate folding actions will be planned in the model fabrication process. Note, in the origami design (e.g. [12]), all the 3D structures are required to be built by folding a single 2D origami sheet; in comparison, in the AAM process, only portion of geometric features will be built by the folding mechanism.

Other geometric features that are hard to be folded will still be fabricated by the layer-based AM process.A 3D model can be decomposed in many different ways for analyzing foldable portions, e.g. by the cellular decomposition, or by direct analysis of a polygonal model. For a given geometry, some portions of a polyhedron may be easily unfolded into a 2D origami structure with small approximation errors. However, it may be difficult to unfold some other geometric features. For example, in a simple test case as shown in Figure 4, the green portion of the object may be easily fabricated through the folding operation; however, the gray portions on the top and bottom of the green cube are difficult to be unfolded. Instead the two cylinders can easily be built using the layer-based fabrication approach.

Instead the two cylinders can easily be built using the layer-based fabrication approach. Hence a detection algorithm is required to decompose a given 3D model into foldable and unfoldable regions.(a) Using layer-fabrication approach to build the bottom cylinder and the unfolded structure(b) Fold the fabricated sheetDesigned model(d) Using the layer-fabrication approach to continue building the top cylinder(c) Folded structureFigure 4: An example of the AAM process.The feature-based model decomposition approaches have been studied before. However, such approaches are generally knowledge-based and not general. Based on the decomposition of the CAD model as shown in Figure 4a, the AMM fabrication process for the model is shown in Figure 4b-4d.

Based on the decomposition of the CAD model as shown in Figure 4a, the AMM fabrication process for the model is shown in Figure 4b-4d. In general, the AAM process needs four main steps as shown in Figure 5: (1) An approximation algorithm is used to convert a given CAD model into a foldable model representation; (2) an unfolding algorithm is used to unfold the approximated model into a 2D origami sheet structure.

Appropriate crease structures are required to be added in the origami structure as well; (3) the 2D origami sheet structure is fabricated and folded into a 3D object; and (4) a post-processing step is applied to the folded structure to bond the creases and to add additional structures if needed.Designed ModelApproximation AlgorithmApproximated ModelPost- ProcessingUnfolding Algorithm3D StructureFolding2D Origami SheetFigure 5: Main steps of the AAM process.Two critical issues in the AAM process include: (i) how to convert a thin-shell structure into a foldable 2D structure; and (2) how to fold a 2D structure into a designed 3D shape. Our study on the issues is discussed in the paper and the remainder of the paper is organized as follows. An unfolding method based on the cellular and CSG decomposition is discussed in Section 3. A foldable structure design is presented in Section 4. The folding mechanism including the angle control is presented in Section 5.

A foldable structure design is presented in Section 4. The folding mechanism including the angle control is presented in Section 5. Some experimental results of the AAM process are presented in Section 6. A fabrication speed comparison between the AAM and traditional layer-based AM processes is also given in Section 6. Finally conclusions are drawn in Section 7 with some future work.3 UNFOLDING OF A CAD MODELA given 3D model can have infinite number of solutions when unfolding it into a 2D sheet. A general decomposition and folding method is desired for a given CAD model. A cellular-based decomposition method is developed for the AAM process including two main steps, (1) approximating a 3D model with a set of voxels, and (2) the unfolding of the voxels into a 2D sheet.3.1 Approximation of a 3D ModelFolding a 2D sheet can only generate a polyhedron with flat surfaces. To better utilize the folding process, it is desired to simplify a given 3D model by removing curved surfaces.

To better utilize the folding process, it is desired to simplify a given 3D model by removing curved surfaces. In addition, a simple geometric representation is preferred in defining and analyzing the given geometry.To unfold a given CAD model, an approximation step is first applied to generate a simpler representation of the 3D model. Two choices have been considered in such an approximation step. The first approximation method is the voxelization by converting a given 3D model into a set of voxels. The second approximation method is based on the Constructive Solid Geometry (CSG) representation by converting a given 3D model into the Boolean results of some CSG primitives [13]. Both approximation methods can lead to a simpler model that is easier to be analyzed. In comparison, (1) the voxelization provides a simple and general approach for any given geometry. The accuracy of the method is related to the resolution of the voxels.

In comparison, (1) the voxelization provides a simple and general approach for any given geometry. The accuracy of the method is related to the resolution of the voxels. The higher the resolution, the better approximation the generated voxels will have. However, the related unfolding and folding process will be more complex. (2) The CSG approximation can use less primitive to capture the original geometry. Consequently, the related folding process can be simpler. However, the decomposition of a complex geometry may require significant efforts since it is difficult to have a general method. Both decomposition methods have been investigated for the AAM process.

However, the decomposition of a complex geometry may require significant efforts since it is difficult to have a general method. Both decomposition methods have been investigated for the AAM process. They are discussed as follows.1121d1c1b1a2b 2c2c22b2b112122a1d11121c1b1d32b1a1b1b11a1b1b211a1a1b22c2b11c1d1d11d21c1d1d31b11b1c1b21e1e1d11d21d1d31f1f1e1fFigure 6: An unfolding algorithm for a voxel-based model.3.2 Unfolding algorithms for voxels and CSG modelsBased on the two aforementioned approximation methods, the unfolding of a 3D model can beclassified as the voxel-based and the CSG-based approaches.(1) For a voxel-based model, several algorithms can be used to unfold the voxel representation into a 2D origami sheet. Figure 6 shows a simple example. Suppose a 3D model is approximated by four voxels, which are used to define the input geometry.

Figure 6 shows a simple example. Suppose a 3D model is approximated by four voxels, which are used to define the input geometry. Note that a voxel model has only flat surfaces with fixed folding angles. Consequently the unfolding problem for voxels can be simplified. As shown in Figure 6, (1) suppose the bottom surface is set as the 2D plane in the unfolding process. (2) The bottom units are labeled based on their columns: the case as shown in Figure 6 has only two columns: 1 and 2. (3) The faces on each column are labeled in a specified orientation as a loop: for example, the faces in column 1 are labeled in the clockwise orientation as 1a, 1b, 1c, etc. The faces in column 2 are labeled in a reverse orientation as 2a, 2b, etc. (4) For each labeled column face, its neighboring faces in the same column that have not been labeled will be marked as its child faces: for example, the child faces of 1d are marked as 1d1, 1d2, and 1d3.

In addition, the child face 1d3 is linked to the face 2b in column 2. (5) The faces of one column can be unfolded in the sequence based on their labels. (6) The child faces of each column facecan be unfolded and added to its neighbors. (7) Similarly, the next column can be processed until all the columns have been unfolded.(2) For a CSG-based model, each basic primitive can be unfolded separately. The unfolded structures can then be assembled based the Boolean operations that are performed between the primitives of the CSG model. Figure 7 shows an example. The designed model can be decomposed into two primitives, a cube and a triangular prism. The unfolded structures of the two primitives are shown in the figure. The Boolean operation between the two primitives is a Union operation. Based on it, a final origami sheet can be assembled based on the topological relations of the two unfolded surfaces (i.e.

The Boolean operation between the two primitives is a Union operation. Based on it, a final origami sheet can be assembled based on the topological relations of the two unfolded surfaces (i.e. faces 1 and 3 are neighboring faces).Figure 7: An unfolding algorithm for a CSG-based model.4 FOLDING OF 2D ORIGAMI SHEETAppropriate crease structures need to be designed in order for a 2D origami sheet to be foldable. The MIP-SL process can easily fabricate multi-layer structures with complex geometry in each layer. In addition, multiple types of materials can be added in the MIP-SL process [14]. In our study, three types of crease structures have been tested.

In addition, multiple types of materials can be added in the MIP-SL process [14]. In our study, three types of crease structures have been tested. One is based on a multi-material crease design, and the other two structures are based on different geometric designs using a single material.4.1 Foldable crease designsFor a 2D origami sheet that is unfolded from a 3D model, the designed creases can have either a softmaterial or a specially designed geometry to make it foldable.(1) To fabricate a foldable 2D structure, a multi-material-based crease design is shown in Figure 8. In the crease model, a soft material is used to enable the folding along the crease while a rigid material is used in the other portions to achieve a better structural property. The fabricated physical model is shown in Figure 8b. The soft portion is shown in transparent, which connects two hard portions in both sides (as shown in yellow). Hence the crease can be folded to a certain angle.

The soft portion is shown in transparent, which connects two hard portions in both sides (as shown in yellow). Hence the crease can be folded to a certain angle. Figure 8c shows a fabricated origami sheet structure by unfolding a box.Figure 8: A multi-material soft crease design and its samples.(2) Two hinge designs based on a rigid material are shown in Figure 9. Figure 9a shows a living hinge structure design. Such a hinge structure has been widely used in various polymer-based lid designs. The hinge can be folded because it has a relatively thin feature in its crease. The relative dimensions of the specific living hinge design will make the structure sufficiently flexible to be folded and unfolded. Another hinge design based on a pattern structure is shown in Figure 9b. The crease portion uses a pre- defined pattern such that the hinge can be flexible and the related structure can be folded and unfolded.

The crease portion uses a pre- defined pattern such that the hinge can be flexible and the related structure can be folded and unfolded. The built physical objects based on the two hinge designs are also shown in the figure.Figure 9: Two foldable hinge designs based on a single material. (a) A living hinge structure design; and (b) a pattern structure based hinge design.4.2 A comparison of the three crease designsBased on our tests, the multi-material-based crease design and the two single-material-based hinge designs can all realize the folding operation of a 2D origami sheet. Table 1 presents the advantages and disadvantages of the three crease designs.For the AAM process, (1) the multi-material-based soft crease has good strength and large bending angle. However, the crease is relatively bulky. In addition, the fabrication process is more complex since two types of materials are required in building the crease model.

However, the crease is relatively bulky. In addition, the fabrication process is more complex since two types of materials are required in building the crease model. (2) The single-material-based living hinge structure is simple, compact, and can achieve a large bending angle as well. However, the structure is relatively weak and easy to be torn due to the thin hinge connection. (3) The pattern structure based hinge model has good strength.

However, the structure is relatively weak and easy to be torn due to the thin hinge connection. (3) The pattern structure based hinge model has good strength. However, the structure is bulky and only has a limited bending angle.Table 1: A comparison of three foldable crease designs.Model typeAdvantagesDisadvantagesSoft creaseGood strength; flexible bending angleBulky, two materialsLiving hingeSimple; one material; Space saving Flexible bending angleWeak connectionPatterned cutsModest strength; one materialBulky, bending angle limitedThe choice of an appropriate crease design depends on the 2D origami sheet design, the folding angle of a crease, and the capability of the AM processes.

For the MIP-SL based AAM process, the single- material-based living hinge design is selected in building the test cases as discussed in Section 6.5 FOLDING ACTUATION AND CONTORLAnother important issue in the AAM process is the folding actuation and the accurate folding angle control. An aforementioned crease design can be folded using either internal or external actuation mechanisms. An internal actuation mechanism means the 2D structure can be folded by some embedded actuators such as a shape memory alloy foil or a shape changeable film. An external actuation mechanism means the structure is folded using some external forces such as robotic hands or some specially designed folding devices. Both internal and external actuation mechanisms have been tested in our study. The test results demonstrate that both approaches can be used in the AAM process.(a)(b)(c)Figure 10: A polystyrene film embedded in a multi-material crease design.

The test results demonstrate that both approaches can be used in the AAM process.(a)(b)(c)Figure 10: A polystyrene film embedded in a multi-material crease design. (a) A PS film based design; (b) a built physical model; and (c) a self-folded structure after heating.5.1 A plastic film based actuation mechanismTo demonstrate the internal actuation approach, a polystyrene (PS) plastic film is embedded in the designed crease structures. The selected PS film is thermally sensitive and will quickly shrink over 30% when temperature overpasses a transition temperature (~120oC). A multi-material soft crease with an embedded PS film is shown in Figure 10. The film is constrained by the rigid material in both sides of the crease. In addition, a soft material is used to cover the bottom of the PS film. Hence the PS film is onlyexposed to the outside environment from the top side.

In addition, a soft material is used to cover the bottom of the PS film. Hence the PS film is onlyexposed to the outside environment from the top side. Consequently, when such a crease is exposed to a heated environment, the top side will shrink while the bottom side is constrained by the soft material. The one-side shrinkage of the PS film will generate a bending movement of the designed crease. A fabricated crease structure based on the MIP-SL process is shown in Figure 10b. After putting the 2D origami structure in a pre-heated oven with a temperature over 120oC, the crease will be bended to a folding angle around 90o in less than 10 seconds (refer to Figure 10c).The PS film can also be embedded in a single-material-based structure design as shown in Figure 11a. The PS film is indicated in green color. It is sandwiched by rigid material (in yellow) in both sides.

The PS film is indicated in green color. It is sandwiched by rigid material (in yellow) in both sides. However, in the crease portion, only one side of the film is covered by the rigid material while the other side is exposed to the outside environment. When the structure is heated, the one-side shrinkage of the PS film will bend the crease towards the side that is exposed to the heated air. Figure 11b shows a fabricated origami sheet structure for an open box model. When the structure is put in a pre-heated oven with a temperature over 120oC, the structure will fold into a structure as shown in Figure 11c in around 10 seconds. The detail analysis of the bending effect and related parameter settings for desired bending angles will be presented in other publications.(a)(b)(c)Figure 11: A polystyrene film embedded in a single-material crease design.

(a) A PS film based design; (b) a built physical model – the embedded PS film is highlighted in red; and (c) a self-folded structure after heating.5.2 An external actuation mechanismAn external actuation approach based on manual operations has been tested. Manual operations, although being slow, can easily fold complex crease structures. In the future, robotic hands or specially designed actuation devices can be developed to replace the manual operations such that the folding process can be automatic. The development of such actuation devices needs to consider three main issues including: (1) the separation of a folding portion from its constraints; (2) determining the holding position of a crease during the folding process; and (3) designing required folding motions that have no collisions.

For a single crease, the development of such an actuation device is straightforward; however, the actuation devices can be complex when multiple hinges in a 2D origami sheet need to be folded by following a certain sequence. The development of such actuation devices for the AAM process will be studied in our future work.5.3 An angle lock designTo achieve a controlled bending angle, an angle lock design has been developed for the living hinge design. As shown in Figure 12, two angle fixtures are added in the both sides of the crease. Based on them, the bending angle can be accurately controlled through their geometrical relationship. In addition, the friction between the angle fixture 1 and 2 can fix the folded structure as well.A folded structureFigure 12: An angle lock designFigure 13 shows a built test case with the angle lock design. The structure is manually folded and could be firmly fixed through the friction between the two angle fixtures.

The structure is manually folded and could be firmly fixed through the friction between the two angle fixtures. Two or more angle fixtures can also be added in the crease to ensure the folded structure to be securely fixed in the designed angle.Figure 13: A built structure with angle lock design.After the built 2D origami sheets are folded into 3D structures, the last step in the AAM process is the post-processing of the folded creases. In this optional step, additional liquid resin can be added to the creases and then be cured by external light sources such as a projector or a portable LED light [15]. Consequently, the two sides of the creases can be permanently bonded by the added cured resin.6 TEST CASESTwo types of test cases are designed for physical experiments including folding during and after the layer-based fabrication process. In the first type, three test cases are designed, which are shown in Figures 14 and 15.

In the first type, three test cases are designed, which are shown in Figures 14 and 15. In the tests, only certain portions of the models are fabricated by the folding approach; the other portions are fabricated by the layer-based AM approach. In the second type, another three test cases are designed, which are shown in Figures 16 and 17. In the tests, the folding operations are used after the whole origami structures have been fabricated by the MIP-SL process. In addition, the fabrication speeds of the AAM and layer-based AM processes are compared for the six test cases.7.1 Test cases using a hybrid fabrication approachFigure 14a shows a designed CAD model to be fabricated by the AAM process. The input model is decomposed into two portions, a foldable thin-shell structure and a top lid. Accordingly the building process based on the decomposed portions is shown in Figure 14b. An unfolded 2D sheet with designed creases based on living hinge structures is fabricated.

Accordingly the building process based on the decomposed portions is shown in Figure 14b. An unfolded 2D sheet with designed creases based on living hinge structures is fabricated. The portions that need to be folded are detached from the platform while the other portions remain on the platform. The folding of the detached 2D structures is realized through manual operations. After the 2D structure is folded into the desired 3D structure, the MIP-SL process will resume until the remaining top layers are finished. Finally the built object will be taken out from the platform for cleaning. Figure 14b shows the built object using the AAM process.Based on a similar building process, foldable structures with more complex features can be fabricated. Two test cases with additional features on the side surfaces are shown in Figure 15.

Two test cases with additional features on the side surfaces are shown in Figure 15. Note that such features can be easily built in the XY plane while it is difficult to fabricate them in the Z axis due to the issues such as over-curing in the Z axis and additional supports in the MIP-SL process. The added supports may also be inside a hollow object and hence will not be easily removed. In comparison, suchfeatures on the side surfaces can be unfolded to the bottom surface in the AAM process (refer to Figure 15b). After the 2D surfaces in the XY plane have been built, the side surfaces can be folded into a 3D structure. Figure 15c and 15d show the fabricated 3D objects.Figure 14: Fabrication procedure of the AAM process.Figure 15: Two test models.Note that the USC-Mickey model as shown in Figure 15d is hollow with a fully enclosed volume.

Other manufacturing processes will not be able to build such a hollow object since tools will not have access paths to its internal features.7.2 Test cases using folding after AM fabricationFigure 16 shows a simple test case based on the folding operations after the MIP-SL process. For a given box model, a foldable origami sheet is designed with the living hinges added to all the creases. The unfolded 2D structure is fabricated. It is then manually folded into a 3D structure. The folding operations need to be controlled to avoid breaking the hinge structures. A tape can be used in temporally fixing the folded structure. In the test, a small amount of liquid resin is added to the folded creases. The resin is then cured by a hand-held light source. After removing the tape, the fabricated box is also shown in Figure 16. The top lid can easily be opened and closed.Figure 16: A test case of an open box model.Figure 17 shows two more test cases.

The top lid can easily be opened and closed.Figure 16: A test case of an open box model.Figure 17 shows two more test cases. Similar to the test cases as shown in Figure 15, some additional features can be added to the side surfaces. Such features can easily be built in the XY plane by the MIP- SL process. The built features can then be folded into planes with desired angles, e.g. 90o in Figure 17a and 60o in Figure 17b.Figure 17: Two test cases. (a) A hollow box model; and (b) a triangular prism model.7.3 Fabrication speed analysisThe AAM process can reduce the fabrication time by converting 3D features that requires the Z axis fabrication process into an unfolded 2D features that requires the fabrication of a single XY plane. Hence the required number of layers can be significantly reduced.Table 2 show a fabrication speed analysis of the six aforementioned test cases for the AAM process. The manual operation time has been considered in the building time analysis.

The manual operation time has been considered in the building time analysis. The folding of the creases will typically take less than 2 to 3 minutes depending on the operator’s experience. For a comparison, the fabrication time of the layer-based AM processes is also shown in the table. In the analysis, it is assumed that the AM process will take 30 seconds in building a layer. Thus the layer number will directly affect the fabrication time. As shown in the table, the AAM process can significantly reduce the fabrication time for the tested thin-shell structures.Table 2: A comparison between the AAM and AM processes on the fabrication speed.7 CONCLUSION AND FUTURE WORKIn the paper, an origami inspired additive manufacturing process named Assembled Additive Manufacturing (AAM) has been developed. The main idea of the AAM process is to integrate the folding based fabrication process with the layer-based fabrication process.

The main idea of the AAM process is to integrate the folding based fabrication process with the layer-based fabrication process. Some general issues related to the AAM process are discussed including the approximation of a CAD model, unfolding of a 3D model, the design of foldable structures and related folding mechanisms. Experimental tests of the AAM process have been performed to compare the fabrication speeds of the AAM and the layer-based AM processes. Some potential benefits of the AAM process have been demonstrated for hollow objects including shortening the fabrication time and removing the needs of supports.A significant amount of work remains in order to mature the newly developed AAM process.

sessing Energy Requirements and Material Flows of Selective Laser Sintering of Nylon PartsCassandra Telenko and Carolyn Conner Seepersad Mechanical Engineering Department The University of Texas at Austin Austin, Texas 78712ABSTRACTSelective laser sintering (SLS) is a prominent technology for rapid manufacturing (RM) of functional parts. SLS and competitive RM technologies are generally assumed to be more environmentally sustainable than conventional manufacturing methods because the additive process minimizes tooling, material waste, and chemical fluids. A thorough life cycle analysis (LCA) of the environmental impacts of SLS has yet to be published. This study focuses on a section of the SLS part life-cycle. It tracks the nylon powder material flows from the extraction and synthesis of the material to SLS part production. Basic material properties and environmental effects are reported.

It tracks the nylon powder material flows from the extraction and synthesis of the material to SLS part production. Basic material properties and environmental effects are reported. Estimates of material waste and energy use are also reported and compared with those of injection molding.MOTIVATIONSLS is a laser-based, additive manufacturing method for creating functional parts. Parts are built from a CAD model, layer-by-layer. The part is built inside a powder bed, and most of this powder can be reused in subsequent builds. The SFF community has identified multiple sustainable advantages of SLS [1]. For example, SLS and other SFF machines reduce transportation and supply chain impacts, as designs can be sent to machines via CAD files and produced on the spot with no special tooling. Quantitative support for many of these hypotheses is not publicly available, and many environmental metrics, such as material waste, toxicity, and process energy use, are not well-known [2].

Quantitative support for many of these hypotheses is not publicly available, and many environmental metrics, such as material waste, toxicity, and process energy use, are not well-known [2]. A 2009 NSF-sponsored workshop identified multiple research needs relating to sustainability, including material performance data and measures of process sustainability in comparison with other manufacturing methods [1]. The research presented in this paper contributes to these needs by providing estimates of the nylon material waste and energy use of SLS in comparison with injection molding (IM).These estimates will contribute to broader life cycle inventories (LCIs) of SLS and a better understanding of how SLS compares with alternative manufacturing processes.

An LCI is an important part of life cycle analysis (LCA), an accepted method for quantifying the environmental impacts of a product or process throughout its life cycle, starting with the procurement of materials and ending with the return of materials to the environment or processing plant [3]. LCI data report energy and material inputs and air, water, and soil emissions and solid waste outputs for each process within the life cycle.In this research nylon and energy LCI data are collected and interpreted for two stages of a part life cycle: material refining and part manufacture. The energy required to create nylon was included in this study to better evaluate the cost of material waste in the SLS and IM manufacturing processes. Additionally, the energy requirements of SLS and IM manufacturingprocesses were considered.

Additionally, the energy requirements of SLS and IM manufacturingprocesses were considered. A full LCI would include additional aspects of these processes, such as chemical inputs and emissions.SCOPEA few studies provide data on material and energy use during SLS, but the data are too few to calculate SLS impacts with any certainty. The scope of this research was limited to the available literature and interviews with local service bureaus. Calculations were performed for a section of the SLS part life-cycle; specifically, nylon powder material and energy ﬂows were tracked from the production of the material through SLS part production. Other life cycle stages were not considered. As injection molding (IM) is the biggest competitor for nylon parts, a similar study was performed for IM, and energy values were compared.

Other life cycle stages were not considered. As injection molding (IM) is the biggest competitor for nylon parts, a similar study was performed for IM, and energy values were compared. Metal SLS parts, which are believed to be more sustainable [1], were not considered.CoalReused PowderCrude OilNatural GasNylon Polymer- izationPowder FormationDuraform PASelective Laser SinteringFinal Part (1kg)Material WasteCoalCrude OilNatural GasNylon Polymer- izationGranulate FormationNylon GranulateInjection MoldingFinal Part (1kg)Material WasteFigure 1: Process Diagrams and Scope of This StudyFigure 1 depicts the gate to gate process being studied. Important parameters are the material yields of the IM and SLS processes and the powder formation process. Electricity is not shown in the flowchart, but was estimated for all processes as well. The following sub-sections describe each of the processes and their possible environmental impacts.

Electricity is not shown in the flowchart, but was estimated for all processes as well. The following sub-sections describe each of the processes and their possible environmental impacts. These process descriptions are followed by a synthesis of the material and energy use quantities from the literature and a calculation comparing IM and SLS yields.PROCESS DESCRIPTIONS AND PARAMETERSNylonNylons, or polyamides (PA), are manufactured from feedstocks of coal, crude oil, and natural gas. Nylon 6-6 and Nylon 6 are the most common types of PA for IM, but Nylon 12 is most commonly used in SLS. Nylon 6 is created from a single monomer, caprolactam. The carpolactam is most commonly created using a hydrolytic polymerization process, described in Kirk-Othmer [4], using a series of products such as benzene and phenol, that are created from coal, crude oil, and natural gas feedstocks.

The creation of laurolactam, the monomer used for nylon 12, is similar to that of caprolactam.The Plastics Europe [5,6] life-cycle inventories for nylon 6 granulate production were used to approximate the energy embodied in a kilogram of nylon 12. Embodied energy is the energy required to produce a product and differs from the energy content of that product. The available data did not distinguish between fossil fuels used as feedstock for the monomer and fossil fuels290290used as energy for processing the feedstock. To determine the value of the fossil fuel energy, it was assumed that nylon 12 and its fossil fuel feedstocks contained equivalent amounts of carbon, and that every kilogram of nylon 12 required 0.7 kg of carbon in its fossil fuel feedstock.

The remaining mass of fossil fuel inputs were thus assumed to be burned during nylon refining.The final estimate for energy use during nylon manufacture was approximated as 85 MJ per kilogram of nylon produced, using a mix of electricity (5.5 MJ/kg) and the energy stored in fossil fuels (79.5 MJ/kg). Output data suggest that material waste from nylon production is much less than one percent and therefore negligible. This assumption is supported by Kirk-Othmer’s discussion of environmental aspects of nylon processing [4].As depicted in Figure 1, the forms of nylon feedstock used in IM and SLS differ. IM uses nylon granulate, and SLS uses a commercial nylon 12 powder that is created using a precipitation process. As described in US Patent 4,334,056, a mixture of polymerized lauryllactam and caprolactam are dissolved in ethanol, reacted, and crystallized with a temperature- and pressure-controlled process [7,8].

Duraform PA, a common nylon 12 powder used in SLS, contains particles that range in size from 25-92 microns for reliable sintering. The patent suggests that only 2% of particles yielded by this process are outside of the acceptable size range. The patents were not used to estimate the energy use of nylon powder production because industrial systems are generally more efficient than single experiments. It is possible that the value is similar to that of metal powder; creation of metal powder is reported to be about 1MJ/kg [9]. Without energy data for nylon powder production, it was assumed that the 85 MJ/kg value included the creation of granulate and powder (i.e. as the dashed boundary in Figure 1.)Although nylon production has harmful environmental releases such as nitrogen compounds, nylon itself is not considered a harmful material. With the exception of additives, nylons pose no significant health risks.

With the exception of additives, nylons pose no significant health risks. There may be hazards from heating the monomer; the MSDS sheet for Duraform PA, a commonly used nylon 12 powder, in Appendix A, indicates that the powder should not be heated above 340 Celsius or combined with acids and oxidizing agents. Nylon does not, currently, biodegrade; methods for reformulating nylon to assist degrading organisms have been successful, but not economical [4]. Recycling of nylon is not yet economically viable, as reprocessing often degrades the material. SLS nylon powders slowly experience this thermal degradation as well and are not infinitely reusable in most commercially available forms.Selective Laser SinteringSLS, shown in Figure 2, builds parts directly from CAD files using a high-powered laser and a bed of material, usually nylon 12, in powder form. Referring to Figure 2, the SLS machine starts by heating the powder to near melting point.

Referring to Figure 2, the SLS machine starts by heating the powder to near melting point. A roller spreads a thin layer of the powder from the delivery bin across the fabrication bed. The laser sinters the layer selectively to match the CAD file’s cross-section. The layer-based sintering process continues until the part build is completed. Un-sintered powders provide support for the sintered parts. After the build finishes, the bed is cooled, and the part is removed. Excess powder, or part cake, is separated from the finished parts and reused.291291Figure 2: Selective Laser Sintering Machine Schematic [10]In theory the powder bed, or support, can be entirely reused, but best practice requires that each part bed contain a fraction of virgin material. As mentioned in the previous section, the powders degrade from repeated heating and cooling. For Duraform PA, the manufacturer recommends a 30% virgin mix [11].

As mentioned in the previous section, the powders degrade from repeated heating and cooling. For Duraform PA, the manufacturer recommends a 30% virgin mix [11]. Specialized nylon powders, such as glass-filled nylon, might require a mix of 40% virgin powder.The quantity of powder wasted by SLS operations is not clear. Dotchev and Yussof [11] report an average build density of 12.68% by weight; for every 100kg of powder that experience the thermal SLS process, only 12.68 kg of part are created on average. They tested ten different builds and recorded a range of powder utilization rates from 3-30%. From these experiments they reported that 80-90% of the powder can be reused.These estimates of material waste were compared with those of a major service bureau. The service bureau reported a 10-15% build density with 10% of part cake (8.5% of total build) lost during the breakout process [12].

The service bureau reported a 10-15% build density with 10% of part cake (8.5% of total build) lost during the breakout process [12]. The service bureau also estimated that its operations scrap about 40% of excess powder, but that other shops might vary from 1-75%.A numerical model was used to analyze powder utilization for ten consecutive builds and determine a possible range of yields. Parameters from both sources were tested with 10% loss of part cake and an idealized build density of 28%. These calculations indicated that up to 44% of the material that enters the SLS process might be wasted and that SLS has a likely range of yields from 56-80%.In addition to powder utilization, it is important to account for energy utilization in the SLS process. Four studies [13-16] exist that report similar power draws, but different specific energy consumptions for the production of SLS nylon parts.

Four studies [13-16] exist that report similar power draws, but different specific energy consumptions for the production of SLS nylon parts. The studies measured power draws ranging from three to twelve kilowatts while operating the EOS EOSINT P760, EOSINT M250 Xtend, SLS Vanguard HiQ+HS, and DTM Sinterstation 2500. Specific energy consumption estimates by these sources ranged from 52 MJ/kg to greater than 140 MJ/kg. This energy range stems partly from variations in build density and height. Mognol et al. [14] built the same part in multiple orientations and found that difference in build height corresponded to a range of 115- 187 MJ. The energy values and build heights reported were not proportional, but the delay required to prepare each layer significantly increased the time and energy required to build the part [14]. It is difficult to prescribe a specific energy consumption constant for SLS, because of the variance in build density and height. Sreenivasan et al.

It is difficult to prescribe a specific energy consumption constant for SLS, because of the variance in build density and height. Sreenivasan et al. [15] reported values that suggest 52.2 MJ/kg can be expected for a realistic part build, and this value was used for the purposes of this paper.292Injection MoldingInjection molding is a common process for producing parts from thermoplastics. The thermoplastic is fed through a hopper into a heated barrel, where it is fed through a nozzle into the cavity of a mold. Pressure is maintained as the material is cooled. After the set cooling time is reached, the finished mold is ejected.Material waste and part rejects can be a problem for some IM operations. According to the PlasticsEurope data sets in SimaPro and GaBi, IM incurs only a 0.5-0.6% material loss [5,6]. Nevertheless, Olmstead cites a “standard reject rate” of 5% [17]. Michaeli and Greif [18] cite that waste content of material can be 5-50%, presumably by weight.

Nevertheless, Olmstead cites a “standard reject rate” of 5% [17]. Michaeli and Greif [18] cite that waste content of material can be 5-50%, presumably by weight. Waste varies by shop and is created during machine start-up, purging, and maintenance. Most IM material can be reused, but the average material loss from these processes is not clear. For this study, a range of material losses, from 1-50% were tested. Energy use for the process was assumed to be 9.9 MJ/kg based on the Plastics Europe databases [5,6]. It is important to note that this value is lower than the averages, 12-19 MJ/kg, reported by Thieriez and Gutowski [19], who utilized environmental impact data from over 100 sources for electrical, hybrid, and hydraulic IM machines.

Nevertheless, the Plastics Europe database is a common source for life cycle analysis data regarding refining and manufacture of plastics.ENERGY AND MATERIAL USE OF SLS AND IMEnergy and material waste estimates were used to compare the efficiency of SLS with IM. Three types of variables were collected. The process characterization, k, describes the amount of energy expended in megajoules per kilogram of product output. The variable, W, indicates mass. The process characterization, Y or yield, indicates the ratio of mass of final product to material feedstock entering the process. The values of these variables are defined in Table 1.Table 1: Process Characterization ValuesSpecific Energy Consumption (MJ/kg)NylonSLSIMYield (output/input)PowderSLSIMPart Weight (kg) All8552.29.90.980.56-0.8 0.5-0.991Two equations were used to calculate the specific energy embodied in an SLS or IM part.

Equations 1 and 2 define the energy required to create a part from IM and SLS, respectively. Both equations are the sums of the energy to make the part during the injection molding process and the energy to make the feedstock entering that process.

(cid:2783)(cid:2785) (cid:4666)(cid:2781)(cid:4667)Equation 4 is a sample calculation, based on Equation 3, using the baseline data recorded in Table 1. The yield for IM is assumed to be the worst case scenario, 50%, in this calculation. Using these values and energy equations, it is expected that an SLS process meet a yield of 68% to be competitive with IM from an energy standpoint.The plot in Figure 3 illustrates a number of yields calculated using equation 3. The darker shaded region encompasses the range of possible IM yields, 50-99%. The lighter region encompasses the range of possible SLS yields, 60-80%. Because of the high energy use of the SLS process, five times the energy use of IM, the SLS process is rarely competitive. The feasible range of values requires that IM have a yield below 60% to be competitive with SLS.

The feasible range of values requires that IM have a yield below 60% to be competitive with SLS. Even if SLS powder were infinitely recyclable (with a yield rate near 1), SLS cannot compete with the standard reject rate of 5% for IM.Yield Rates for Equal Embodied Energyn oi t c a r Fd e i y S L Sl2 1.8 1.6 1.4 1.2 1 0.8 0.6SLSIMFeasible Range0.40.50.60.70.80.91Injection Molding (IM) Yield FractionFigure 3: SLS is Only Competitive with Inefficient IM Operations294Breakdown of Embodied Energy for 1 kg part made using SLS at 60% and IM at 95% Yield2009.4MJ18016052.2 MJ140Nylon Electricity120MFG ElectricityJ M100806083.7 MJ5.6 MJ4.8 MJ5.1 MJ135.2 MJMFG Fossil FuelNylon Fossil Fuel40200IMSLSFigure 4: A larger percentage of SLS energy use is electrical energySLS is less competitive than IM primarily because it uses more powder and electricity during the manufacturing process.

Figure 4 provides a breakdown of embodied energy for representative yields. Overall, the SLS part requires 196.8 MJ while the IM part requires about half of that amount, 99.2 MJ. SLS requires about 1.6 times the energy of IM for nylon production, and 5 times the energy of IM for manufacture. The total embodied energy estimate for an IM part is similar to that reported by Thieriez and Gutowski [17].It is important to note that these values do not include maintenance and infrastructure costs. Die production, for example, is important in IM operations, but absent from SLS manufacture. The costs of creating and maintaining these machines at difference production volumes also affects their specific energy consumptions. Loughborough University’s ATKINS project feasibility study includes a discussion of possible logistical impacts and values [20].

Loughborough University’s ATKINS project feasibility study includes a discussion of possible logistical impacts and values [20]. For example, a cost study by Hopkinson and Dickens [21,22] at Loughborough University found that SLS is cheaper than IM at production volumes less than 12,000 units. These cost estimates indicate how the larger supply chain, labor and maintenance costs compare and could indicate their energy requirements as well. Morrow et al. [9] compared the environmental impacts of creating tools and dies from machining and Direct Metal Deposition (DMD) and found that about 8 MJ are required to machine a single die, but did not include the extra material costs incurred by this special tooling. It is not obvious how the scaling of SLS and tradition IM operations change their environmental impacts.CLOSUREThis research quantified the material and energy use of SLS nylon parts and compared these estimates with IM parts.

A survey of the literature found loose agreement for the amount of nylon powder wasted during SLS processing and few distinct values for nylon material waste from IM. For these reasons, a variety of possible material yields were calculated and compared.The results indicate that SLS nylon parts are not as efficient as IM parts when considering nylon material and energy consumed during the material and part production process. It is possible that these findings extend from energy consumption to pollution emissions295as well. For example, Thieriez and Gutowski report that most IM emissions occur during material production [19]. Even if the SLS process has fewer on-site chemical emissions than IM, it is probable that emissions from greater nylon manufacture offset this advantage.Future work could extend this research to a number of areas. Supply chain effects such as reduced transportation and infrastructure costs could make SLS more favorable.

Supply chain effects such as reduced transportation and infrastructure costs could make SLS more favorable. These effects could be studied on a large scale with assumptions about the number and locations of available SLS machines. More research comparing metal SLS processing with competitive manufacturing processes, such as machining and die-casting, could provide different insights as well.Additionally, one of the advantages of SLS is its ability to produce parts that cannot be manufactured using IM. These parts, with optimized geometries, have the opportunity to significantly increase the efficiency of end-use applications. For example, SLS can create honeycomb structures with optimized strength to weight ratios.

[16] Kellens, K., Dewulf, W., Deprez, W., Yasa, E. and Duflou, J.R., “Environmental Analysis of SLM and SLS Manufacturing Processes”, Proceedings of 2010 CIRP International Conference on Life Cycle296Engineering, Hefei, China, May 19-22 2010 pp. 423-428 [17] Olmsted B., and Davis M., 2001, Practical Injection Molding, Marcel Dekker, Inc., NY, NY. [18] Michaeli W., and Greif H., Training in Injection Molding: A Text and Workbook, Hanser Gardner Publications, Munich. [19] Thiriez, A. and T. Gutowski. “An Environmental Analysis of Injection Molding," IEEE International Symposium on Electronics and the Environment, San Francisco, California, USA, May 8- 11, 2006.

ssessing new support minimizing strategies for the additive manufacturing technology SLM”M. Cloots*, A.B. Spierings*, K. Wegener†Inspire-institute for rapid product development irpd, Lerchenfeldstrasse 5, 9014 St.Gallen, Switzerland † ETH Zurich, Institute for machine tools and manufacturing IWF, Tannenstrasse, 8092 ZurichAccepted August 16th 2013AbstractTo successfully produce metal parts by SLM, additional support structures are needed to support overhanging surfaces in order to dissipate process heat and to minimize geometrical distortions induced by internal stresses. These structures are often massive and require additional post-processing time for their removal. A minimization of the extent to which support structures are needed would therefore significantly reduce manufacturing and finishing efforts and costs. A specific component segmentation strategy is developed.

A specific component segmentation strategy is developed. It allows the segmentation of critical areas of the component by applying a specific scanning strategy with appropriate energy input and optimized supporting strategies. The results indicate that the supporting effort can generally be reduced, e.g. overhang geometries with an angle to the horizontal of less than 35° can be manufactured without any support. The successful realization of the segmentation strategy in combination with optimized support structures allows the implementation of a stacking strategy, thereby using the available work space more efficiently.IntroductionSelective Laser Melting (SLM) belongs to the additive manufacturing process, which is characterized by the layerwise production of the component. The high degree of geometric freedom, which is a unique feature of the additive processes, is often limited by process-specific effects.

The high degree of geometric freedom, which is a unique feature of the additive processes, is often limited by process-specific effects. In the SLM process, high temperature gradients, and consequently high internal stresses are the cause for the limitation in the degree of complexity (Pohl H., 2001). In SLM research there are the numerous attempts dealing with the fundamental development of strategies to avoid of tension and deformation in the component. Kruth points out that the exposure strategy has a significant influence on the stress development within the component (Kruth et al., 2004). From this, recommendations for a stress-optimized powder exposure are formulated. Mumtaz shows a completely new approach in the use of eutectic alloys (Mumtaz, 2011). In comparison to conventional metallic alloys, eutectic alloys have a low melting point. Therefore, due to lower required process temperatures, the thermo-mechanical effects decreases.

In comparison to conventional metallic alloys, eutectic alloys have a low melting point. Therefore, due to lower required process temperatures, the thermo-mechanical effects decreases. Another way to counteract component thermomechanically induced defects is demonstrated by Ott and Buchbinder (Zaeh & Ott, 2011) (Buchbinder, Schilling, Meiners, Pirch, & Wissenbach, 2011). By heating up the base plate, a decrease in the temperature gradients can be achieved. It is proven that defects in the microstructure and strains in the part can be significantly reduced. The successful conversion and use of the described strategies leads automatically to a diminished utilization of support structures. A segmentation strategy is introduced in this work as another instrument to avoid massive support structures. With the segmentation strategy, the component is segmented according to its position in the construction field in critical and uncritical regions.

With the segmentation strategy, the component is segmented according to its position in the construction field in critical and uncritical regions. Uncritical are those regions ofthe workpiece which are surrounded with sufficient solid material underneath able to act as a sufficient heat sink dissipating the process heat. Regions of the workpiece are critical when sufficient heat dissipating is unavailable.

Regions of the workpiece are critical when sufficient heat dissipating is unavailable. Overhanging structures and surfaces are typically classified as critical, because of their limited possibilities to conduct process heat in lower regions of the workpiece.In this work, two objectives are pursued:1) Supported overhangProcess parameters are to be determined for a maximum overhang of 0 °, which should be designed with a minimal support structure.2) Unsupported overhangThrough the use of appropriate process parameters and exposure strategies, the smallest possible critical overhang angle is presented, in which the support is omitted.Figure 1: Example of a segmented sample used to describe the objectives of this workA further benefit of these two goals is to make apparent the possibilities of using stacking strategies.

The stacking strategy helps to make the SLM process more economical, due to using the available space in the SLM-machine more efficiently.Experimental setup and methodologySLM-MachineThe investigations are performed on an industrially prepared system: Concept Laser M2. It nm includes The beam has a M² of 1.3 and a measured diameter of 90 microns (D4σ) in the focal plane. The focused beam is measured using a Spiricon SP620U beam analysis camera.a200 W Nd:YAGfiberlaserata wavelengthof1070Powder used A series of experiments are carried out with 316L stainless steel powder. The powder fraction has a particle size distribution of D10= 7.12 µm to D90 = 24.17 µm.Surface quality The surface quality is measured tactilely using a perthometer by Mahr.Density measurement The Archimedes method is used to determine the sample density (Spierings A.B., 2011).Minimum Support For the studies, a special support structure is developed.

The support structure used here has the following properties.1) The connection to the part is realized only via fixation points. The fixation points have the task to connect the part and dissipate the process heat. The purely selective connection of the component on the support intends to simplify and to ease the separation of the support structure from the component. 2) Due to the lattice structure, there are no walled structures, which imbeds powder. So the waste of powder can be avoided. 3) The grid structure is used for support rigidity so that horizontal and vertical forces can be absorbed by the crossbars.Figure 2: Realized (left) and schematic (right) of support elementsMethodology The results are analyzed using Design of Experiments (DoE).

The answer Y (the measured value) will be expressed according to the process parameters to the following relation (factors: see Table 1) to the following relation (development of Taylor):Y = a0 +a1X1 + a2X2 + a3X3 +a12X1X2 + a13X1X3 + a23X2X3 + a123X1X2X3 2 2 + a33X3 +a11X12 + a22X2constant 1st order interactions 2nd orderWhere X1, X2 and X3 are the used influence factors and ai the coefficients of the model. The determination of the various coefficients ai of the models is carried out by the least square method.sample [no.] 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16normalized process parameter X3 X2 X1 -1 1 1 1 -1 -1 1 0 0 -1 1 -1 0 0 -1 -1 0 0 0 0 1 0 0 0 0 -1 0 0 1 0 0 0 0 1 -1 1 1 1 1 -1 -1 -1 1 1 -1 -1 -1 1Table 1: Central Composite Design (CCD) planFigure 3: Graphical representation of CCD planPairwise comparison Pairwise comparison is a tool to rank a set of decision-making criteria and rate the criteria on a relative scale of quality.

In this work the quality is ranked from 1 (=fine) to 16 (=poor).Experiments, Part 1: Supported OverhangProcess parameters and support density for critical layersWhile applying the DOE to a selected overhang geometry needing support structures, an optimization has to be found for the following three goals: high dimensional stability, surface quality and material density. The support used here is characterized only by selective attachment sites at the critical layers. Consequently, a moderate energy input in the production of critical layers has to be chosen. Due to the low heat dissipation of such support structures (Figure 2), a high energy input would result in unintended agglomerations of powder particles.Figure 4: principle of connection layers and support structures634Therefore, it is necessary to develop process parameters, which guarantee a minimum of support and a good connection to the densest possible material for the critical layers above the support.

After finishing the so-called connection layers, these serve as a heat sink for the subsequent standard layers. To ensure the production of the part using –“standard layers” (Figure 4) on the connection layers, it is important to investigate which process parameters generate defect-free connection layers. In parallel, it has to be investigated how large the maximum possible distance between the fixation points of the minimum support should be in order to build the connection layers successfully.

In parallel, it has to be investigated how large the maximum possible distance between the fixation points of the minimum support should be in order to build the connection layers successfully. In Table 2, the DOE factors used are listed.Process parameter Laser power Layer thickness Exposure strategy X1 = hatch distance X2 = distance fixation point X3 = scan speed Table 2: Process parameters for supported overhang experimentstype constantunit[W] [µm] [-] [mm] [mm] [mm/s]variable200 30 unidirectional 0.08/ 0.095 / 0.11 2 / 4 / 6 1,000 / 1,500 / 2,000The combination of the connection layers and the minimum support substitutes the use of solid support structures to a wide degree and contributes to a more efficient post-processing.

To determine the optimal SLM process parameters (scan speed and hatch distance) for the connection layers in the critical area, 24 x 24 mm² samples consisting of 5 connection layers, which are based on different densely populated minimum supports (distance fixation point), are produced. 5 connection layers are sufficient to evaluate the influence of the process parameters. Within the study 16 samples are created such that experiments are designed using a Central Composite Design experiment plan.Figure 5: Sample CompositionThe production of the 5 connection layers are realized with a unidirectional hatch strategy, thereby minimizing heat concentration effects at the end of the single scan lines. Preliminary investigations show that the more economical bidirectional hatch strategy does not produce satisfactory surface structures at the connection layers. The reason is the inhomogeneous energy input of the bidirectional hatch strategy.

The reason is the inhomogeneous energy input of the bidirectional hatch strategy. When evaluating the samples to determine the best process parameters set for the production of connection layers, their surfaces are qualitatively analyzed with regard to a closed defect-free surface (figure 8, sample 12 and 14) which could influence the ongoing process.Quantitative measurements of the surface have been omitted. Nevertheless, to represent a gradation in the results obtained, a pairwise comparison is chosen for evaluating the surfaces and defining the surface quality via a rank position (1 = fine; 16 = poor).Number of necessary layers in the critical region In a further step, it is examined how many connection layers are actually needed to provide sufficient thermal and mechanical support for the subsequent layers under normal production conditions.

These samples are generated, which consist of minimal support, connection layers and, if necessary, standard layers.Figure 6: Sample composition for optimal number of connection layersThe first three connection layers are always generated with a unidirectional area scan strategy. The remaining layers are produced with island scanning (Spierings, 2009), using a unidirectional scanning, to minimize the induced stresses and strains.type process parameter # connection layers variable Table 3: Number of connection layers within the experimentsunit [-]20 / 60 / 100Experiments, Part 2: Unsupported OverhangConventionally, the construction of overhanging structures with an angle to the horizontal below a certain value is only possible through the use of support structures. The primary objective of support structures is to dissipate process heat and shrinkage.

The primary objective of support structures is to dissipate process heat and shrinkage. Since there is powder underneath the outer contour of the overhang, the process heat can only be dissipated laterally, as metallic powder material has a comparably low thermal conductivity (Sih & Barlow, 2004). At decreasing overhang angles, the difficulties in heat dissipation arise. Overhang angles smaller than 45 ° can negatively affect the surface quality due to heat accumulation (Wang, Yang, Yi, & Su, 2013) and consequently adhering metallic powder particles. By combining certain exposure strategies with low heat input process parameters giving a lower heat input, it is possible to reduce the overhang angle without any losses to surface quality and work piece geometry.

Thereby, the influence of the hatch distance, scan speed and scan angle are investigated.The scanning strategy will be varied in that way that the laser exposure runs parallel to the contour / edge of the component (scan angle 0 °), runs perpendicular to the edge of the component (scan angle 90 °) or a combination of both (scan angle 45 °) (Table 4). For all parameter combinations, the exposure of a layer always starts inside the sample and ends at the edge. The exposure begins therefore always inside the sample, to prevent excessive energy input at the edge of the sample where heat dissipation is difficult.

The visual impression of the surface quality and the quality achieved at the edges contribute to the overall impression of the sample. Since the overhang areas represent a final, and thus visible surface in contrast to the connection layers, the surface quality of the overhanging faces are also measured quantitatively for the evaluation of the parameters. In addition, the generated component density is also determined.Results, Part 1: Supported OverhangThe distance between the fixation points has the greatest influence on the results. As expected, the best results are achieved when the minimum distance (2 mm) is used. The remaining parameters scan speed and hatch distance have less influence on the final results.

As expected, the best results are achieved when the minimum distance (2 mm) is used. The remaining parameters scan speed and hatch distance have less influence on the final results. This is particularly evident with samples 12 and 14.(sample # = parameter set #)Figure 7: Main effects on surface quality (left); visualized model (right)Figure 8: Resulting track energy; surface quality and density; surface examples of fine and poor resultsBoth quality. However, the energy input for sample 14 is almost 3 higher times as the energy input of sample 12. Both parameters will be used as a benchmark for further investigations.samplesshowexcellentresults, withregardstosurfaceNecessary number of connection layers When determining the optimal number of connection layers, parameter set of sample 12 and 14 are used.

With both parameters, it is, compared to other combinations, possible to realize good surface qualities, thereby generating optimal conditions for building up subsequent layers.The influence of the volume energy becomes apparent when studying the optimal number of connection layers. The lattice structure of the support cannot avoid the resulting shrinkage of the connection layers. This is particularly evident in the samples, which are produced with parameter set 14. Thereby, the samples created by using parameter set 12 show a much lower shrinkage, compared to the samples prepared with parameter 14. Furthermore, the investigations show that one can already continue the build process using standard parameters after only 20 connection layers. Up to here it is possible to say that the connection to the minimal structure, in principle, is possible.

This is particularly noticeable in the separation of the component from the support. Samples, which were produced with parameter set 12, can be separated from the support without sample. no mechanical Samples, which were produced with parameter set 14, can be separated only mechanically (e.g. sawing) from the support.parametertheusingas whenstrongsetnotiscomponentasTherethesupportrestsassistance.onareResults, Part 2: Unsupported OverhangFigure 11: Main effects on surface quality (left); visualized model (right)With all 16 parameter combinations, the samples are built successfully up to an overhang angle of 30 °. In the test series with an overhang angle of 25 ° not all process combinations perform to high qualitative sample. Four samples of the test series are of such a weak quality that a measurement of the sample surface on the overhanging side is not possible (Figure 13, sample 1).

Four samples of the test series are of such a weak quality that a measurement of the sample surface on the overhanging side is not possible (Figure 13, sample 1). Since the quality of the results varies strongly within the 25 ° test series, the samples of this test series are ideally suited to determine the influence of process parameters. Figure 11 (left) shows the results of the qualitative assessment, which are transferred into the Taylor series model previously mentioned. Thus, the scan angle has the greatest influence on the sample quality such that scan angles of 0° lead to the best results. Regarding the scanning speed, it can be said that with high scanning speeds better results are typically achieved. The measurement results of the tactile surface measurements underline once again the result of the qualitative evaluation samples. For scan angles of 0°, the best surface results are achieved across all overhang angles (Figure 12).

For scan angles of 0°, the best surface results are achieved across all overhang angles (Figure 12). The achieved surface qualities, Ra, lie between 13 and 20 µm. Not only the best results are obtained with the scan angle of 0 °, the smallest investigated overhang angle of 20° is only possible with a scan angle of 0° (Figure 13). The achievable sample densities are between 86.9% and 97.2%.Figure 12: surface qualities (scan angle; overhang angle)Figure 13: track energy; surface quality and density; sample 2 (fine) and sample 1 (poor)Results, Part 3: Stacking StrategyThe application and refinement of the chosen process parameters for supported and unsupported overhang situations can make the SLM process even more effecient. Until now, the number of parts placed on a base plate is limited by the lateral dimensions of the plate.

Until now, the number of parts placed on a base plate is limited by the lateral dimensions of the plate. With the stacking strategy, a new dimension is realized and allows increasing the number of manufacturable parts and improved machine utilization.Figure 14: Demonstrator samplesThe examples in figure 14 show that this strategy is feasible in principle.ConclusionThe segmentation strategy is derived from the fact that the heat dissipating conditions during manufacturing are inhomogeneous. For this reason, components are segmented into critical and non-critical areas, which are manufactured with specific process parameters and exposure strategies. In this work, therefore, two particularly challenging component situations were investigated.

In this work, therefore, two particularly challenging component situations were investigated. At the maximum possible overhang of 0 ° (supported situation), it was shown that good results are possible with the use of minimal support given that the fixation points are at a distance of 2 mm, and can be separated with little effort from the support structure. Already after 20 connection layers, subsequent standard layers can be produced. Depending on the energy input and accuracy, a part density between 89.4% and 98.4% can be reached. The results of the overhang tests (unsupported situation) show that an overhang angle of 20° is realizable. The best results in surface quality and part density are obtained by using a scan angle of 0°. The maximum possible part density at the smallest possible overhang of 20° is 95.1%. The surface qualities Ra are in a range of 13 to 20 µm.

viewed PaperAssessing the feasibility of process window estimation for L-PBF Ti-6Al-4V through hardness measurementsArun Poudel, Mohammad Salman Yasin, Shuai Shao, Nima Shamsaei*National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849, USA Department of Mechanical Engineering, Auburn University, Auburn, AL 36849, USACorresponding author: shamsaei@auburn.edu Phone: (334) 844-4839AbstractOptimizing the process window for an additively manufactured material using X-ray computed tomography (XCT) is a time-, labor-, and capital-intensive process where a large number of coupons need to be analyzed in the process parameter space. Hence, as a cheaper and faster alternative, this study assessed the relationship between relative density and hardness of laser powder bed fused (L-PBF) Ti-6Al-4V. Coupons with different levels of porosity were fabricated by changing the recommended laser power in the range of ±20% in an interval of 10%.

Coupons with different levels of porosity were fabricated by changing the recommended laser power in the range of ±20% in an interval of 10%. The martensitic microstructure didn’t vary significantly with the change in laser power. On the other hand, Rockwell C hardness (HRC) did. HRC was observed to be in good correlation with the defect content in the coupons. Higher defect content in the coupons resulted in lower HRC and vice versa. Hence, the hardness measurement technique can be used to estimate the process window of L-PBF Ti-6Al-4V.Keywords: Laser powder bed fusion (L-PBF), Non-destructive testing, X-ray computed tomography, Process window, Rockwell hardnessIntroductionOne of the major distinctions that separate the additively manufactured (AM) parts from the conventionally manufactured ones is the presence of a higher number of volumetric defects [1,2].

During loading, these defects act as stress risers and can deteriorate both tensile properties as well as fatigue performance of AM parts [3–6]. Generally, post-fabrication treatment such as hot isostatic pressing (HIP) is implemented to reduce the degree of porosity and its influence on the mechanical performance [7–12]. However, the complete removal of entrapped pores, as well as large-sized defects from AM parts, is difficult [4,13–15] if not impossible. Another way to improve the mechanical properties, especially fatigue, of AM parts is to minimize their formation during the AM process.To minimize defect formation, AM parts are fabricated utilizing a predetermined set of process parameters such as laser power, laser speed, hatch distance, etc., residing well within the process window for a specific material and AM process [16]. Typically, a process window is established based on the relative density of the parts fabricated using different process parameters [17–20].

Typically, a process window is established based on the relative density of the parts fabricated using different process parameters [17–20]. Relative density is typically measured using different techniques such as Archimedes, X- ray computed tomography (XCT), micrography of cross sections, etc. [21,22]. Among these1895techniques, the high-resolution XCT has been observed to estimate the density and characterize the defects in the AM parts with good accuracy [22,23].The XCT is an advanced imaging technique that utilizes an X-ray beam to detect the voids and foreign particles in a body/part [24]. Even though XCT is an accurate nondestructive evaluation (NDE) technique, its results are highly dependent on the voxel size used during the scans. The use of different voxel sizes has been observed to result in different defect information even for the same part [25]. Furthermore, performing an XCT scan and post processing the data is a time-and capital-intensive process.

Furthermore, performing an XCT scan and post processing the data is a time-and capital-intensive process. The emergence of new AM materials as well as techniques demands a fast and relatively robust approach to establish the process window. In this regard, considering the microstructure of the material is constant, the hardness has been observed to correlate well with the porosity level [26,27]. Hence, hardness measurements can be considered an appealing alternative to assess the porosity and estimate the process window for AM materials [28,29].Few studies have been done to investigate the hardness of AM materials fabricated using different process parameters [26,27]. However, a thorough investigation correlating the porosity with the hardness of the AM material is still lacking. Hence, this study fabricated the coupons by altering the laser power from manufacturer recommended process parameters and obtained the process window via both XCT and Rockwell C hardness (HRC) measurements.

This assessed the feasibility of process window estimation for laser powder bed fused (L-PBF) Ti-6Al-4V Grade 5 material using the hardness measurements.Material and methodsPlasma atomized Ti-6Al-4V Grade 5 powder supplied by Advanced Powders and Coating (AP&C) with a size distribution in the range of 15 µm to 53 µm was used to fabricate coupons using EOS M290, a laser powder bed fusion (L-PBF) AM machine. The chemical composition of the feedstock powder as provided by the manufacturer is shown in Table 1. Different level of porosity was induced by altering the recommended laser power from +20% to -20% in the intervals of 10% (i.e., P+20%, P+10%, P0%, P-10%, and P-20%). The coupons were named Px% which suggests that the laser power was changed by x% during its fabrication with all other process parameters remaining constant.

The coupons were named Px% which suggests that the laser power was changed by x% during its fabrication with all other process parameters remaining constant. The EOS recommended process parameters for Ti-6Al-4V Grade 5 material include 280 W laser power, 1200 mm/sec laser speed, 30 µm layer thickness, 140 µm hatch distance, 67° inter-layer rotation, and 10 mm stripe width. The geometry and dimensions of the coupons used for XCT and hardness tests are shown in Fig. 1. The XCT scans were performed using ZEISS Xradia 620 Versa machine and hardness tests using the LECO LCR-500 HRC tester.Table. 1 Chemical composition of the Ti-6Al-4V grade 5 powder provided by AP&C [30]. OAlVFeCNHTi <0.001 Bal.YWt. % 6.260 3.910 0.210 0.140 0.010 0.010 0.0021896Fig. 1 Geometry and dimension of L-PBF Ti-6Al-4V coupon in (a) as-built and (b) machined conditions.The hardness tests were performed in accordance with ASTM E18-20 [31]. For hardness tests, the as-built coupons (see Fig.

For hardness tests, the as-built coupons (see Fig. 1(a)) were mounted in a cold set epoxy resin and polished using sandpapers with grit sizes ranging from P280 to P360. Furthermore, it was polished from both sides to make the coupons flat as well as to let hardness coupons touch the tester platform. Eight indents were made using a diamond spheroconical indenter in the transverse plane; i.e., perpendicular to the build direction in the top cylindrical portion of every coupon. According to ASTM E18-20, a 150kgf total test force was applied for each indentation. The microstructural analysis of these coupons was performed using Zeiss 550 scanning electron microscope equipped with a back-scatter electron (BSE) detector. Similar to the hardness tests, the BSE imaging was performed on the transverse plane. Regarding XCT, the coupons were machined to a geometry seen in Fig. 1(b). The smaller-sized coupon enables the high-resolution scans (i.e., voxel size of 1 µm) in its center.

Regarding XCT, the coupons were machined to a geometry seen in Fig. 1(b). The smaller-sized coupon enables the high-resolution scans (i.e., voxel size of 1 µm) in its center. The scans were performed using 160 kV voltage and 25 W power X-ray source. After the scan, the reconstructed absorption data were post-processed using Dragonfly and ImageJ software [32,33] to calculate the relative density of the fabricated coupons.Results and discussionProcess window obtained via XCTDifferent level of porosity was induced in coupons fabricated by altering the laser power as seen in Fig. 2. The underheated (i.e., laser power lower than recommended) coupons mostly consisted of irregularly shaped lack of fusions (LoFs), while overheated (i.e., laser power higher than recommended) ones had small-sized gas-entrapped pores (GEPs). The size and number of defects in the underheated coupons increased with a decrease in the laser power during the fabrication process.

The size and number of defects in the underheated coupons increased with a decrease in the laser power during the fabrication process. In the case of P0% coupon, small-sized spherical defects and one irregular- shaped defect, shown using a green arrow in Fig. 2 were observed. The levels of porosity in recommended and overheated ones seemed to be similar. No keyholes (KHs) were observed in the overheated coupons. The volume of the defects observed in Fig. 2 was quantified and the relative density of the coupons is presented in Fig. 3. According to the observation made using Fig. 2, the relative density of the underheated coupons decreased with decreasing energy inputs during the fabrication process. In addition, the relative densities of the recommended and overheated coupons1897were observed to be approximately equal.

In addition, the relative densities of the recommended and overheated coupons1897were observed to be approximately equal. It was attributed to the sufficient melt-pool overlaps in between the adjacent laser track as well as layers—prevents the formation of LoFs—and also, not enough energy to induce KHs.Fig. 2 Visualization of defects induced in coupons fabricated using different laser power. The green arrow points toward an irregular-shaped defect, most likely an LoF in P0% coupon. The scale bar corresponds to all coupons.Fig. 3 Relative density of coupons fabricated using different laser power.Process window obtained via hardness measurementsThe HRC of the Ti-6Al-4V coupons fabricated by different laser power is shown using a box plot in Fig. 4. The HRC was observed to follow a similar trend as the part density of the coupons as seen in Fig. 3. The average HRC of the recommended coupon was the highest among the considered coupons.

4. The HRC was observed to follow a similar trend as the part density of the coupons as seen in Fig. 3. The average HRC of the recommended coupon was the highest among the considered coupons. In underheated coupons (i.e., P-10% and P-20%), the mean of the HRC decreased and its range increased with the decrease in the energy input. In other words, the HRC values decreased with the increase in the defect content in the underheated coupons. Similarly, in the case of overheated coupons (i.e., P+10% and P+20%), the absence of defects led to similar HRC values as recommended ones. Fig. 4 illustrates a significant overlap in the range of HRC between recommended and overheated coupons. It is in agreement with the density measurements performed via XCT.1898Fig. 4 Comparison of HRC among the coupons fabricated using altered laser power.

It is in agreement with the density measurements performed via XCT.1898Fig. 4 Comparison of HRC among the coupons fabricated using altered laser power. In the legend, IQR denotes the interquartile range.In order to understand the role of microstructure on the hardness of the fabricated coupons, BSE imaging was performed and the resulting images are shown in Fig. 5. The martensitic microstructure was observed in all the coupons. The lath thickness did not vary significantly with the variation in the laser power used during fabrication. The hardness of Ti-6Al-4V material is dependent on the grain size [27]. The hardness value decreases with the increase in the grain size and vice versa. The similarity in the microstructure of the coupons fabricated using different laser power (see Fig. 5) suggests its minimal influence on the variation of their hardness values as shown in Fig. 4. Hence, the variation in the hardness values was attributed to the defect content in the coupons.

5) suggests its minimal influence on the variation of their hardness values as shown in Fig. 4. Hence, the variation in the hardness values was attributed to the defect content in the coupons. As seen in Fig. 2, the increase in the defect content in the underheated coupons resulted in their lower hardness. In addition, the increase in the number and size of the defects in P-20% coupon as compared to P-10% has increased the range of variation in their hardness values. On the other hand, the overheated and P0% coupons have similar hardness values due to the similar level of defect content (see Fig. 2).1899Fig. 5 BSE images of L-PBF Ti-6Al-4V coupons fabricated using different laser power.ConclusionsThe feasibility of the hardness measurement technique to correlate the defect contents as well as to estimate the process window for L-PBF Ti-6Al-4V material was studied. Throughout this study, the following conclusions were drawn:1.

Throughout this study, the following conclusions were drawn:1. Hardness was in good correlation with the defect content in the material. The hardness was observed to decrease with an increase in the defect content and vice versa.2. The rate of decrease in the hardness was higher for underheated coupons as compared to the overheated ones for the same percentage change in the laser power.3. The martensitic microstructure of the as-built L-PBF Ti-6Al-4V didn’t vary significantly with the change in the laser power.AcknowledgmentsThis material is based upon the work partially supported by the National Science Foundation (NSF) under grant # 1919818 and the Federal Aviation Administration (FAA) under grant # FAA- 12-C-AM-AU-A2.References[1] A. Yadollahi, N. Shamsaei, Additive manufacturing of fatigue resistant materials: 14–31. Int.Challenges https://doi.org/10.1016/J.IJFATIGUE.2017.01.001. S. Lee, N. Ahmad, K. Corriveau, C. Himel, D.F. Silva, N.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAssessing the Impact of the Powder Production Method on Ceramic-filled Polyamide Composites made by Laser SinteringI. Kletetzka*, F. Neitzel*, H.-J. Schmid*Direct Manufacturing Research Center (DMRC) and Particle Technology Group, Paderborn University, GermanyAbstractPolymer composites represent the industry standard in injection molding for the production of plastic components with increased requirements in terms of heat resistance and stiffness. In the field of laser sintering (LS), these materials are less common so far. In order to extend the available material variety for the LS process, new ceramic-filled Polyamide 613 powders are investigated within the scope of this work. Here, the resulting properties from two different powder production methods are compared.

Here, the resulting properties from two different powder production methods are compared. One filled powder is produced by dry blending and the other powder with the same filler and filling ratio is produced by encapsulating the filler particles inside the polymer particles within the dissolution-precipitation process. It was found that encapsulating the filler particles can provide certain benefits for the processability, for example an improved powder flowability or better filler dispersion. However, encapsulating the filler also alters the thermal properties of the precipitated powder.IntroductionLaser sintering (LS) is an additive manufacturing technology that uses a laser to selectively fuse small polymer particles into solid 3D objects. To do this, the polymer powder is deposited in thin layers on the build platform, preheated by IR heaters to just below the melting temperature, and finally fused by the laser according to the component cross-section.

Then the build platform is lowered by one layer height, which is typically around 100 µm. These four steps are repeated until the parts are finished [1–3]. Laser sintering has evolved from a rapid prototyping technology to a process for the direct production of complex plastic components, since it allows the production of complex geometries without the need for support structures and provides a good combination of productivity and detail resolution compared to most other additive manufacturing technologies for thermoplastic materials [3, 4].For producing functional prototypes or final components, the functional and mechanical properties of additively manufactured parts have become increasingly important [5, 6]. However, the limited selection of LS materials currently available restricts the ability to meet all customer- specific requirements.

However, the limited selection of LS materials currently available restricts the ability to meet all customer- specific requirements. A survey carried out by Sculpteo reveals that more than 50 % of the participants think that 3D printing needs a wider selection of materials to keep growing (see figure 1). When selecting a material for additive manufacturing 72 % of the participants consider the mechanical properties of the material to be very important. Furthermore, a low material price (52 %) and easy to use materials (44 %) are important to users of additive manufacturing.

Furthermore, a low material price (52 %) and easy to use materials (44 %) are important to users of additive manufacturing. [7]181What does the 3D printing industry need to grow?More SoftwareTraining and EducationEasier to operate TechnologiesSustainable Technologies and MaterialsLower Cost of EntryNew MaterialsMore Reliable Technologies0%20%40%60%80%Percentage of respondentsFigure 1: Survey on the influencing factors which limit the growth of 3D printing [7]Material properties of plastic components can be influenced not only by using different polymers but also by adding fillers to the polymer to create a compound. In the injection molding industry, filled polymers represent the industry standard for components with increased requirements regarding heat resistance and component stiffness. Filled or reinforced polymers are therefore often used in the automotive industry [8].

Filled or reinforced polymers are therefore often used in the automotive industry [8]. By adding fillers to laser sintering polymers, the part stiffness, heat resistance and abrasion resistance can potentially be improved while reducing the material cost at the same time [9–12]. Furthermore, flame retardancy or conductivity of polymers can be altered by adding functional fillers [13–15]. Despite these potential advantages, filled materials are rarely used in laser sintering compared to injection molding.The aim of this paper is to assess the impact of the powder production method on the processability of ceramic-sphere-filled polyamide 613 powder and on the resulting composite properties made by laser sintering. While ceramic-filled LS materials have shown promise in improving mechanical properties [16], their use in industry is still limited. Within the study a dry blended powder will be compared to a powder where the filler particles are encapsulated within the polymer particles.

Within the study a dry blended powder will be compared to a powder where the filler particles are encapsulated within the polymer particles. The overall objective is to increase the range of available materials for laser sintering and to create a better understanding of the impact of different production methods of polymer composite powder.Powder Production MethodsIn order to produce polymer powders with fillers for the LS process, there are essentially three different methods suggested by researchers in this field. These are dry blending, melt mixing and encapsulating as illustrated by figure 2.Dry blending is the fastest and easiest way to produce filled powders for laser sintering. In this process, a polymer powder is mixed with filler particles, e.g. with a drum hoop mixer or a rotary mixer. Various authors use mixing times in the range of 15 min to 24 h [17–19].

In this process, a polymer powder is mixed with filler particles, e.g. with a drum hoop mixer or a rotary mixer. Various authors use mixing times in the range of 15 min to 24 h [17–19]. Dry blending182requires both the polymer and the filler to already be in powder form with a size suitable for laser sintering (preferably about 20 - 80 µm [1, 18, 20]). Depending on the particle size, particle morphology and densities of the polymer and filler, it may be difficult to achieve a homogeneous mixture [21]. After the mixing process, the powder is typically left in the feed tanks of the LS system for at least a few hours to allow the electrostatic charges generated during mixing to dissipate. The advantage of dry blending is that commercially available plastic powder can be quickly mixed with various fillers and filling ratios, allowing the material to be customized for specific build jobs.

A possible disadvantage is that agglomeration or segregation effects may occur depending on the differences in particle size and density between filler and polymer particles. In addition, the adhesion of filler particles and the polymer matrix is based solely on pressure and not on coalescence in the build process [22].Dry BlendingMelt MixingEncapsulatingFigure 2: Powder production methods for filled laser sintering materials183Compounding and subsequent cryogenic grinding is another option for the production of filled polymer powders. This process route is referred to as melt mixing in figure 2. In this process, the starting materials do not have to be in powder form. The polymer granules or powder are compounded with the filler in a twin-screw extruder. The resulting compound pellets are then cooled with liquid nitrogen and ground to powder.

The polymer granules or powder are compounded with the filler in a twin-screw extruder. The resulting compound pellets are then cooled with liquid nitrogen and ground to powder. Since grinding usually results in a particle size distribution that is too broad, the coarse particles are separated by sieving or other classification methods and returned to the mill. In some cases, it is also necessary to remove fine particles from the final product, as these affect the flowability of the powder. The milled powder usually has an irregular shape with sharp edges, which limits the flowability of the powder and therefore negatively impacts the processability [23, 24]. Further rounding steps can improve the flowability of the powder [3]. The mechanisms of particle rounding can be divided into thermal, thermomechanical, mechanical, and chemical, depending on the mode of action.

The mechanisms of particle rounding can be divided into thermal, thermomechanical, mechanical, and chemical, depending on the mode of action. The advantages of this manufacturing process lie, on the one hand, in the flexibility with regard to the materials to be processed and, on the other hand, in the fact that the resulting particles usually contain both filler and plastic. This prevents segregation and may reduce health concerns associated with loose short fibers. In addition, there could be advantages in terms of homogeneity of distribution of the particles in the matrix and in terms of coalescence [25].The method used in this work to produce particles with incorporated filler particles is to coat the filler particles with a polymer coating within the precipitation process. Either plastic granules or plastic powder can be used for the precipitation process.

Either plastic granules or plastic powder can be used for the precipitation process. However, the particle size and shape of the filler particles must be within certain limits, since the filler is not ground in the process and the morphology of the filler particles has a considerable influence on the size and shape of the particles produced in the precipitation process. Unfilled standard polyamide 12 powder is usually produced in a precipitation process from ethanolic solution. The PA12 is dissolved in ethanol under pressure at high temperatures (> 140 °C) and the solution is then cooled slowly and in a controlled manner. Precipitation of the PA12 from the oversaturated solution is initiated by withdrawing the organic solvent and cooling the reactor. The particle size depends on the stirring speed in the reactor [3, 26, 27].

The particle size depends on the stirring speed in the reactor [3, 26, 27]. If a dispersed filler material is present in the solvent during the precipitation process, the filler acts as a heterogeneous nucleating agent for the polymer molecules. Therefore, the filler particles are encapsulated in the resulting powder particles. The filler particles have an influence on the crystallization process and can therefore affect the thermal properties of the resulting powder. The advantage of this powder production method is similar to the potential advantages of melt blending: the filler particles are incorporated into the polymer particles. Therefore, a good dispersion of the filler can be achieved and no segregation should occur [28, 29]. In addition, the precipitation process can produce powders with good sphericity and flowability, which are economical for the production of PA12.

In addition, the precipitation process can produce powders with good sphericity and flowability, which are economical for the production of PA12. A disadvantage is the intensive use of solvents and the necessary know-how and careful process control to obtain particles with the desired properties [10].As can be seen from the sources listed, some research work has already been put into the development of polymer powders with fillers. Most publications concern dry blended powders. In particular, a systematic comparison of the powder properties of powders with incorporated filler particles to a dry blend consisting of the same ingredients is missing so far.

In particular, a systematic comparison of the powder properties of powders with incorporated filler particles to a dry blend consisting of the same ingredients is missing so far. Therefore, the aim of this investigation is to compare a dry blended and encapsulated PA613 based LS powder with the same formulation to analyze the advantages and disadvantages of each powder production method.184MaterialsThe materials for this project were provided by Evonik and are based on their VESTOSINT 3D 8754HT1 PA613 powder. Standard, neat PA613 powder is created by a dissolution precipitation process, similar to standard PA12 powder but can provide better heat resistance and mechanical properties compared to PA12 [30]. In order to further increase the stiffness and heat resistance of the material, ceramic spheres with a mean particle size of 30 µm were added as a filler material.

In order to further increase the stiffness and heat resistance of the material, ceramic spheres with a mean particle size of 30 µm were added as a filler material. Two different powders with the same formulation were produced by dry blending and by encapsulating the filler particles inside of the polymer particles in the precipitation process. For both material 35 wt% of ceramic spheres were added to the matrix polymer.The dry blend was created by simply mixing the neat PA613 powder with ceramic spheres in a mechanical mixer. The powder with encapsulated filler particles was produced by using the dissolution precipitation process. First PA613 was dissolved in heated organic solvent. The ceramic beads were dispersed in the solution by stirring. Afterwards the PA613 was precipitated from the solvent.

First PA613 was dissolved in heated organic solvent. The ceramic beads were dispersed in the solution by stirring. Afterwards the PA613 was precipitated from the solvent. Herby, the ceramic filler particles act as heterogenous nuclei and are therefore coated by the polymer.Experimental MethodsAt first, the powder properties of neat PA613 powder, filled dry blended PA613 powder and filled PA613 powder with encapsulated ceramic spheres are investigated. Particle size distribution, morphology, powder flowability and thermal properties are analyzed to compare the suitability of these powder for use in laser sintering.For analyzing the particle morphology, a Phenom World scanning electron microscope was used, which allows a maximum magnification of up to 24,000 times.

In order to create sufficient conductivity of the samples, the samples were sputtered before imaging.A QICPIC cam sizer manufactured by Sympatec was utilized to determine the particle size distribution in compliance with the ISO 13322-2 standard. A RODOS dry dispersion unit and a VIBRI powder feeder were used for dispersion. The measuring range is set to M5, theoretically allowing a detection of particles in the range of 1.8 µm and 3755 μm. [31, 32]The Hausner Ratio as defined in VDI 3405 Sheet 1.1 is used to determine the powder flowability. The Hausner Ratio is defined as the quotient bulk density and tapped density of the powder. Before testing, the powder was deagglomerated and then 250 ml of powder were filled into a measuring cylinder. The measuring cylinder was closed afterwards and tapped uniformly onto a solid surface. The tapping was repeated, until no more compaction and therefore reduction of the powder volume occurred.

The measuring cylinder was closed afterwards and tapped uniformly onto a solid surface. The tapping was repeated, until no more compaction and therefore reduction of the powder volume occurred. The Hausner ratio was calculated by dividing 250 ml by the volume, that the powder had after compacting it by tapping. The test is repeated three times and the mean value and standard deviation are given. [33]In order to ensure that the correct ratio of ceramic spheres was added to the dry blended and encapsulated material and to check if segregation effects occurred, an ash residue analysis according to DIN EN ISO 3451-1 method A is carried out. For this purpose, a small sample of185powder or composite part is taken and placed in a crucible inside a Phoenix muffle furnace from CEM at 625 °C for 60 min.

For this purpose, a small sample of185powder or composite part is taken and placed in a crucible inside a Phoenix muffle furnace from CEM at 625 °C for 60 min. After complete ashing of the organic constituents, which in this case is the matrix polymer, the crucible with the sample is cooled to room temperature and the remaining filler is weight to determine the filling ratio.The thermal processing window of the different powders was analyzed by differential scanning calorimetry (DSC) with a DSC 214 Polyma from Netzsch. In accordance with DIN EN ISO 11357-3, powder samples with a mass of 5 mg were heated and subsequently cooled down in an aluminum sample pan. Herby, the heat flow between the sample and an empty reference pan was monitored to identify the melting and recrystallization temperature of the material. During the investigations, the samples were heated up to 280 °C at a heating and cooling rate of 10 K/min.

During the investigations, the samples were heated up to 280 °C at a heating and cooling rate of 10 K/min. After the first heating and cooling cycle a second cycle was used to compare the materials independent of the thermal history of the different powder production processes. [34]After characterizing the important properties of the different powders, the materials were processed on an EOS P396 laser sintering system. Suitable laser parameters for unfilled PA613 are known from previous investigations [16, 30]. The parameters were adapted for the filled variants by varying the laser power and additionally testing double exposure strategies for optimized coalescence. Suitable laser parameters were identified by analyzing the mechanical properties and part resolution for different parameters.

Suitable laser parameters were identified by analyzing the mechanical properties and part resolution for different parameters. The selected build parameters are given by table 1.Table 1: Build parameters for investigated powders ParameterNeat PA613Dry blended PA613 Encapsulated PA613Layer height [µm]120120120TemperatureBuild chamber [°C]206.5206.5188Removal chamber [°C]140140140Global beam offset [mm]0.320.320.32Hatch (Hatch exposure 1 / Hatch exposure 2)Laser power [W]38 / -26 / 2626 / 26Scan Speed [mm/s]475047504750Hatch distance [mm]0.250.250.25Offset [mm]0.30.30.3Contour (contour line 1 / contour line 2)Laser power [W]24 / 2424 / 2024 / 20Scan speed [mm/s]300030003000Offset [mm]0.15 / 00.15 / 00.15 / 0After determining suitable parameters, test specimens were built from all three powder and characterized. Tensile tests, compression tests and heat resistance tests were carried out.

Tensile tests, compression tests and heat resistance tests were carried out. Additionally, the parts were analyzed by computer tomography (CT).The mechanical properties of polyamides are strongly dependent on the moisture content in the material, therefore a precise adjustment of this value is necessary to obtain repeatable186measurements. As stated by DIN EN ISO 1110, the mechanical properties of polyamides should be tested after performing an accelerated conditioning process to ensure they have reached the equilibrium moisture level, that would occur after several years of use in standard climate conditions. Therefore, specimens were stored in a climate chamber at 70°C and 62 % relative humidity for 168 h. After conditioning in the climate chamber, the specimens were kept at standard climate conditions (23°C 50 % relative humidity) for a minimum of 72 h before mechanical tests were carried out. In addition to the tests in conditioned state, specimens were tested in dry state.

In addition to the tests in conditioned state, specimens were tested in dry state. They were therefore kept in closed plastic bags together with silica pads to prevent moisture absorption before testing. [35]Tensile tests were carried out as described in DIN EN ISO 527. Sample geometry 1A was used and tested on an Instron 5569 universal testing machine with a 5 kN load cell. The test speed for determining the modulus of elasticity is 1 mm/min and is then increased at 0.27 % elongation until fracture. In accordance with the material standard for polyamides, unreinforced samples were tested at 50 mm/min and filled materials are tested at 5 mm/min. For every parameter set, average mechanical properties were calculated using 5 test specimens and standard deviations are given. The pre-load for all tests was 0.1 MPa with activated specimen protection. It is indicated in each case whether the samples were tested in the dry or conditioned state.

The pre-load for all tests was 0.1 MPa with activated specimen protection. It is indicated in each case whether the samples were tested in the dry or conditioned state. [36, 37]The mechanical properties under compressive loads may differ from those under tensile loads. Fillers in particular can cause this effect. Therefore, the modulus for compression loads is determined additionally to the tensile values. The compressive modulus is determined according to DIN EN ISO 604. For the compression tests an INSTRON 5569 universal testing machine with a 5 kN load is used. For the compression modulus measurement, the strain is recorded optically. The test speed is 1 mm/min. Unless otherwise stated, 5 samples are tested at a time and mean and standard deviation are given. It is indicated in each case whether the samples were tested in the dry or conditioned state.

Unless otherwise stated, 5 samples are tested at a time and mean and standard deviation are given. It is indicated in each case whether the samples were tested in the dry or conditioned state. [38]As one of the project goals was to increase the heat resistance of the material, the heat deflection temperature is measured according to DIN EN ISO 75 method A. Test specimens with dimensions 80 × 10 × 4 mm are used in the flat lying test position. The HDT value is calculated as an average from three test specimens and the standard deviation is given with error bars. [39, 40]Powder Characterization and ProcessabilityThe particle size distribution (PSD) and particle morphology dictate the powder flowability and the powder bed density and therefore have a strong impact on the quality of LS parts. Poor flowability can lead to recoating errors and a low powder bed density can decrease mechanical properties.

Poor flowability can lead to recoating errors and a low powder bed density can decrease mechanical properties. The particle size distribution impacts the surface quality and limits the lowest possible layer height within the process as well. Finer particles can improve the surface quality but limit powder flowability. Spherical particles are desirable for best powder flowability and powder bed density.

Finer particles can improve the surface quality but limit powder flowability. Spherical particles are desirable for best powder flowability and powder bed density. [3]187PSD neat PA61343,53) 3 q ( g2,52l1,510,5010xm / µm100Figure 3: SEM image neat PA613 virgin powder (left) and PSD (right)PSD dry blended PA61343,532,5) 3 q ( g2l1,510,5010xm / µm100Figure 4: SEM image PA613 dry blended with ceramic filler (left) and PSD (right)PSD encapsulated PA61343,532,5) 3 q ( g2l1,510,5010xm / µm100Figure 5: SEM image PA613 with encapsulated ceramic filler (left) and PSD (right)188Neat VESTOSINT 3D 8754HT1 PA613 polymer powder has a regular potato shaped morphology, as shown in figure 3 on the left. Some smaller sub particles adhere to the particle surface. The average sphericity of the particles is ØΨ = 0.852.

Some smaller sub particles adhere to the particle surface. The average sphericity of the particles is ØΨ = 0.852. The particle morphology is typical for a precipitation process and seems well suited for processing in laser sintering. The particle size distribution is shown on the right and is typical for laser sintering powder. There are some larger particles in the material, so layer heights below 100 µm would not be advisable. The d10, d50 and d90 values are given in table 2.Figure 3 shows the PA613 powder dry blended with ceramic beads. The polymer particles and the ceramic beads can be clearly distinguished in the SEM image. The ceramic beads have a very high sphericity, a smooth surface and are smaller compared to the PA613 powder (see d50 values in table 2). The mixed powder therefore has a higher proportion of fine particles compared to neat PA613 powder.The particle morphology of the PA613 powder with encapsulated ceramic beads is shown in figure 4.

The mixed powder therefore has a higher proportion of fine particles compared to neat PA613 powder.The particle morphology of the PA613 powder with encapsulated ceramic beads is shown in figure 4. In contrast to the dry blend, the filler particles cannot be seen. This indicates that the ceramic beads are well embedded into the base polymer. The particles have more of an edgy particle shape compared to neat PA613. The sphericity of the encapsulated powder ØΨ = 0.832 is slightly lower compared to the neat PA613 ØΨ = 0.852. Compared to powder produced by cryogenic grinding, the particle morphology is still better suited for laser sintering, since the particles have smooth surfaces, and the sphericity is only slightly lower compared to standard neat PA613. There are a few fine powder particles visible that potentially do not have any ceramic particles embedded inside.

There are a few fine powder particles visible that potentially do not have any ceramic particles embedded inside. The particle size of the powder with encapsulated ceramic beads is similar to neat PA613 and therefore suitable for laser sintering.The actual filling ratios of both powders was determined by performing an ash residue analysis. The results are listed in table 2. The volume averaged outer radius of the encapsulated powder 𝑟𝑎 can be calculated from the volume share of filler 𝑣𝑓 and the radius of the filler particles 𝑟𝑓, assuming a homogeneous filler distribution as well as spherical particles (equation 1).3 𝑟𝑎 = 𝑟𝑓 ∗ √1 𝑣𝑓(1)Using this equation and the powder characterization results, the mean particle size of the encapsulated powder can be calculated as 66.8 µm.

Under the given assumptions, this value shows good agreement with the measured median value of 68.0 µm.Table 2: Comparison of sizes distributions, sphericity and filling ratiod10d50d90Average sphericityFilling ratio(µm)(µm)(µm)ØΨvol%Neat PA613 (unfilled)34,4864,5782,870,8520Dry blended PA613 ceramic filler32,9863,1990,500,8518,25Encapsulated PA613 ceramic filler40,3868,0098,070,8329,52Ceramic filler after ashing15,8730,4852,500,899189Bulk density, tapped density and the Hausner ratio are determined for all three powder variations. The results are given by figure 6. As expected, the bulk density and the tapped density of the filled materials is higher compared to neat PA613 since the density of the ceramic filler (6 g/cm³) and glass (2.5 g/cm³) is higher than that of neat PA613 (1.04 g/cm³).Neat PA613 has a Hausner ratio of around Hr = 1.2.

Powders with a Hausner ratio below 1.25 are classified as having good flowability. Above 1.25, powder flowability is classified as limited and above 1.4 cohesion dominates. The encapsulated material exhibited an even better flowability with Hr = 1.15 compared to neat PA613 even though the particle morphology was a bit more edgy. Since the encapsulated material is a “prototype material”, it cannot be fully ensured that exactly the same amount of flow aid (nanosilica) is present in the material compared to neat PA613. The higher flowability could be due to the use of more flowing aids. To protect Evonik’s intellectual property, the used flow aid cannot be further specified. The Hausner ratio of the equivalent dry blend is Hr = 1.30 and is therefore much higher than that of the encapsulated material. Since the value is above 1.25, the dry blend with ceramic beads is classified as having a reduced flowability.

Since the value is above 1.25, the dry blend with ceramic beads is classified as having a reduced flowability. It is known that a higher portion of fine particles can cause worse powder flowability, however this effect is quite extensive in this case. Previous investigations with filled PA613 powder also revealed a worse powder flowability if finer filler particles were added, but the effect was not quite as extensive [16].0,9bulk densitytapped densityhausner ratio) ³m c / g (0,80,7/ y t i s n e d r e d w o p0,60,50,40,30,20,10neat PA613PA613 encaps. ceramic beads v1 9.5 vol%PA613 dry blend ceramic beads 8.3 vol%Figure 6: Comparison of Hausner ratios of dry blended and encapsulated powderThe thermal properties of neat PA613, dry blended and the encapsulated material were investigated with DSC measurements in order to identify the sintering window and to analyze differences.

The first temperature cycle shows that the neat PA613 has an onset melting temperature of 208.6 °C and a recrystallization onset temperature of 176.4 °C and therefore a wide sintering window (see figure 7). As expected, the dry blended material has a similar melting temperature, since the polymer powder is same, just mixed with fillers. The melting enthalpy is lower since the filler does not melt. During recrystallization of the polymer filler mixture, the filler can act as an heterogenous nuclei and cause recrystallization at higher temperatures [41]. The recrystallisation peak temperature of the dry blend is slightly higher than that of neat PA613.The encapsulated material exhibits different thermal properties compared to the dry blend. It has a secondary melting peak before the main melting peak at a lower temperature.

It has a secondary melting peak before the main melting peak at a lower temperature. The filler can1901,451,41,351,31,251,21,151,11,051o i t a rr e n s u a Hnot only act as a nucleus for crystallization for the polymer melt, but also during the dissolution precipitation process. During precipitation, the polymer molecules interact with the filler particles and therefore coat the surface of the filler. This can cause differences in the crystalline structure of the polymer and therefore influence its thermal properties. The secondary melting peak is not desirable for laser sintering. Further investigations showed that with adapted parameters during the precipitation, this secondary peak can be avoided, but the investigations within this work were carried out with the material that showed the secondary peak. Further publications with the optimized material will be released in the future.

Further publications with the optimized material will be released in the future. The encapsulated material also exhibited a higher recrystallization temperature compared to the dry blend. It is assumed that this effect is caused by a higher amount of silica flowing aids present in the encapsulated powder.Figure 7: DSC of neat PA613, dry blended and encapsulated variants first temperature cycleFigure 8: DSC of neat PA613, dry blended and encapsulated variants second temperature cycle191After the first temperature cycle an identical second temperature cycle is measured in the DSC (figure 8). Since the thermal history is erased by melting the material in the first temperature cycle, the encapsulated and the dry blended material should show similar properties. The melting temperature of all materials is lower than in the first cycle and the melting enthalpy is significantly lower.

The melting temperature of all materials is lower than in the first cycle and the melting enthalpy is significantly lower. This shows that the crystallinity of the laser sintering powder is higher before melting it once and that the cooling rate in the dissolution precipitation process is apparently lower than the 10 K/min used for the DSC measurements. In the second temperature cycle, the encapsulated and the dry blended material show very similar melting temperatures as one would expect.For processing the powders on an EOS P396, the powder bed temperature was set by determining the temperature necessary to prevent curling of the parts. For the encapsulated material, a powder bed temperature of 188°C was necessary. This temperature is higher than the peak temperature of the secondary melting peak (184.3°C) that the encapsulated powder exhibited in the DSC measurements.

This temperature is higher than the peak temperature of the secondary melting peak (184.3°C) that the encapsulated powder exhibited in the DSC measurements. This resulted in very pronounced caking of the powder bed to the point that it was difficult to remove the parts from the powder bed after the build job (figure 9 middle). With some effort, the parts could still be removed from the powder bed and the adhering powder could be removed by glass bead blasting. Surface quality of the parts build with encapsulated powder was good (figure 9 right). For the dry blended material and neat PA613 a powder bed temperature of 206.5 °C was selected to prevent curling. Parts could be built without problems and without severe caking of the powder bed with these powders. Different exposure parameters were tested and the parameters that could provide the best results were selected (table 1).

Different exposure parameters were tested and the parameters that could provide the best results were selected (table 1). The same exposure parameters were used to build parts from dry blended and encapsulated powder.Figure 9: Curling of Specimens at temperatures lower than 188 °C with encapsulated powder (left), pronounced caking of the powder bed with encapsulated powder (middle) and part from encapsulated powder after glass bead blasting (right)One potential disadvantage of dry blending polymers powders and fillers is that segregation effects could occur inside the powder containers during the build job. The powder containers are usually fluidized with compressed air, to ensure optimal flowability of the powder. In the experiments of this work, the powder was flushed with nitrogen every 3 min for 10 s. If there are192major differences in density or size between the polymer powder and the filler, there is a risk that the individual components will segregate.

If there are192major differences in density or size between the polymer powder and the filler, there is a risk that the individual components will segregate. In the LS process, this effect would mean that the filler content in the components would not remain constant over the height of the build job. Thus, constant material properties could not be guaranteed. To measure possible segregation effects, tensile test specimens built at different heights in the same build job are analyzed by burning out the polymer matrix and determining the filling ratio according to the ash residue method. If the filling ratio is constant over the full height of the build jobs, it can be assumed that no segregation effects occurred.

The results for the PA613 powder with encapsulated ceramic spheres, dry blended ceramic spheres and two further PA613 dry blends with glass beads are given in figure 10.% o vl12%% o v /l50%45%/ e u d i s e r h s a o t g n d r o c c a o i t a r g ni10%8%6%4%2%e u d i s e r h s a o t g n d r o c c a o i t a r g ni40%35%30%25%20%15%10%5%i l l i f0%i l l i f0%Part built at the bottomPart built in the middlePart built at the topPart built at the bottomPart built at the topPA613 encapsulated ceramic PA613 dry blend ceramic beadsPA613 dry blended Silibeads d50 =39 µm PA613 dry blended Silibeads d50 = 58 µmFigure 10: Segregation measured by ash residue analyses of parts built at different built heights for dry blended and encapsulated ceramic filler (left) and dry blended glass beads (right)The filling ratio is between 8 and 10 vol% for the dry blended and encapsulated ceramic filled powders.

Surprisingly, the measured variation of the filling ratio was higher for the encapsulated powder compared to the dry blend. The dry blend showed no significant variation in filling ratio. Also, further investigations with glass bead filled dry blend revealed no significant differences in the filling ratio over the build height. It is therefore deduced that no significant segregation effects occur for dry blends in the feed tanks of an EOS P3 system, if the filler particles are in the range of 30 µm to 60 µm. The encapsulated powder shows a slightly reduced filling ratio towards the top of the build job. Since no segregation occurred in the dry blend and the filler is encapsulated by polymer, no segregation effects should theoretically occur for the encapsulated material either. The encapsulated powder was not mixed after the precipitation process. Therefore, inhomogeneities in the filling ratio from the production process could still be present and cause this effect.

The encapsulated powder was not mixed after the precipitation process. Therefore, inhomogeneities in the filling ratio from the production process could still be present and cause this effect. For further investigations, the complete batch of precipitated powder will be mixed before processing to rule out this cause of the inhomogeneities.193For analyzing the homogeneity of the filler distribution within the printed parts, CT scans of laser sintered cubes (10 mm x 10 mm x 10 mm) were used. Exemplary sectional views of these scans can be found in figure 11.Figure 11: Filler distribution in parts built with encapsulated powder (left) and dry blended powder (right)Comparing the scans of the encapsulated and dry blended material, some differences can be found. For the dry blended material (figure 11 right), the filler particles often occur in small agglomerates, where a few ceramic beads are in direct contact with each other.

For the dry blended material (figure 11 right), the filler particles often occur in small agglomerates, where a few ceramic beads are in direct contact with each other. Additionally, regions with a lower concentration of filler particles are visible. Here mainly molten PA613 is present. The part seems to have a low porosity for a laser sintered cube.The ceramic particles in the cube built with the encapsulated material (figure 11 left) seem to be better dispersed. Less agglomerations can be found in the sectional views. However, the part built from the encapsulated material has a much higher porosity (8.9 %) compared to the dry blend (4.9 %). Especially in the middle of the cube of lot of pores can be found. It is assumed that the differences in thermal behavior are the main cause of the porosity. Since the encapsulated material had to be processed at 18.5 K lower part bed temperature, the layers cool down faster after being molten by the laser.

Since the encapsulated material had to be processed at 18.5 K lower part bed temperature, the layers cool down faster after being molten by the laser. On top of that, the encapsulated material has a higher recrystallization temperature, meaning the exposed layers will recrystallize faster. According to the widely used Frenkel-Eshelby model for coalescence of molten particles, the coalescence increases over time the particles are in molten state. The process is faster for particles with a lower melt viscosity and higher surface tension. According to this model the coalescence for the encapsulated powder will194likely be worse and the porosity therefore higher [42–44]. The higher porosity of the parts built from encapsulated powder will most likely have a negative impact on the material properties.Figure 12 shows the filler distribution within dry blends that have been investigated previously.

Here, a higher filling ratio of spherical fillers or the use of glass fibers led to worse agglomerations than observed for the dry blend with ceramic spheres. These investigations showed that the tendency for agglomeration increases with smaller filler particles, higher filling ratios and fillers with higher aspect ratios (fibers). For producing powders with these fillers, encapsulating the filler might provide even bigger benefits regarding the filler matrix distribution compared to the investigated powder with ceramic spheres and a filling ratio of less than 10 vol%.

The observed agglomeration for dry blends hindered the successful processing of the materials in figure 12 and encapsulating could be a potential solution for this problem.Figure 12: Agglomerate of glass beads in part built from PA613 dry blend with 20 vol% glass beads with a d50 of 32 µm (left) and agglomerated glass fibers in dry blended PA613 (right)Part PropertiesTensile specimens according to DIN EN ISO 527 were built with laser parameter specified in table 1 from neat and filled PA613 powder in X, Y and Z direction and tested in dry state and after an accelerated conditioned procedure according to DIN EN ISO 1110. The results are given by figure 13 to 15.The dry blend and the material with encapsulated ceramic spheres both provide a similar increase in modulus of about 50 % in the dry state and after conditioning the specimens compared to neat PA613.

The tensile strength of the filled material is slightly lower than for the neat material, indicating limited filler matrix adhesion. After conditioning, the difference in tensile strength of neat and filled PA613 is even bigger, indicating a further loss of filler matrix adhesion as previously noticed for glass bead filled PA613 [45]. While the dry blend exhibited isotropic strength properties, reduced properties were observed for the encapsulated material in z-direction. This effect is most likely caused by the lower powder bed temperature and the higher porosity of the195encapsulated material.

As expected, the elongation at break of the filled materials is reduced, however the encapsulated material performed again worse in z-direction compared to the dry blend.40003500a P M/ s u u d o Mls ' g n u o Y30002500200015001000Encapsulated PA613Dry blended PA613Neat PA6135000X dryX condY dryY condZ dryZ condFigure 13: Tensile Modulus for ceramic filled and neat PA613 powder6050a P M/ h t g n e r t S el i s n e T40302010Encapsulated PA613Dry blended PA613Neat PA6130X dryX condY dryY condZ dryZ condFigure 14: Tensile strength for ceramic filled and neat PA613 powder3530% / k a e r b t a n o i t a g n o E25201510Encapsulated PA613Dry blended PA613Neat PA613l50Y dry Figure 15: Tensile elongation at break for ceramic filled and neat PA613 powderX dryX condY condZ dryZ cond196Comparing the stress strain curves of the different materials, differences in the linear part of the stress strain curve are noticeable.

Figure 16 shows three exemplary stress strain curves of neat PA613, dry blended PA613 and PA613 with encapsulated ceramic filler in conditioned state. For each curve, the stress level that causes a deviation in elongation of 0.2 % compared to the ideal elastic line has been calculated and is visualized. This value will be called 𝜎0,2 in this report. The deviation from the linear stress strain behavior could theoretically either be caused by slipping of the macro molecules and resulting plastic deformation of the polymer itself or loss in filler matrix adhesion and a resulting reduction in stress in the material at a given elongation. The higher the 𝜎0,2 value is, the higher loads the material can withstand without showing a large deviation from the ideal linear material behavior. For metals, the 𝑅𝑃 0,2 is used if the material has no distinctive yield point in order to calculated the limit of the elastic material behavior [46].

For metals, the 𝑅𝑃 0,2 is used if the material has no distinctive yield point in order to calculated the limit of the elastic material behavior [46]. Since polymers are known to have a time dependent viscoelastic material behavior, the 𝑅𝑃 0,2 value is not defined for polymers. However, the 𝜎0,2, which is defined similarly to the 𝑅𝑃 0,2 for metals within this report, can still give an indication on stress levels that cause a plastic deformation or damage to the material at the given strain rate that was used for the tensile testing. Even though this method has its limitation for polymers, a larger linear area of the stress strain curve is still desirable. The described method has been used to calculate the 𝜎0,2 value for neat, dry blended and encapsulated PA613 material in dry and conditioned state.

The described method has been used to calculate the 𝜎0,2 value for neat, dry blended and encapsulated PA613 material in dry and conditioned state. In addition to the mean value for the five tested tensile specimens, the standard deviation is given by figure 16.505045454040a P M/s s e r t S3530252015a P M/ 2 0 a m g i S,35302520151010550 0,0%1,0%2,0%3,0%4,0%5,0%6,0%7,0%0X dryStrainFigure 16: Stress strain diagram of conditioned specimens printed in X-direction (left) and calculated 𝜎0,2 values for ceramic filled and neat PA613 powder (right)Even though the encapsulated material exhibited very similar tensile strength and stiffness compared to the dry blend in the conditioned state, there are significant differences visible for the 𝜎0,2 values. The encapsulated material has a linear material behavior up to higher tensile stress compared to the dry blend.

The encapsulated material has a linear material behavior up to higher tensile stress compared to the dry blend. This can be beneficial since these parts can withstand higher loads before experiencing permeant deformation or damage to the material. This is an interesting observation indicating that the filler matrix adhesion in the encapsulated material can transfer197X condslightly higher forces before failing. At the maximum tensile strength and beyond that, the material still fails due to loss of filler matrix adhesion as is visible by the lowered tensile strength compared to unfilled material.In addition to the tensile modulus, the compression modulus of the different materials was analyzed (figure 17). It was found, that for compression loads the spherical filler provided an even higher increase in part stiffness. In Z-direction both the dry blended and the encapsulated powder have a higher modulus compared to the XY-plane.

In Z-direction both the dry blended and the encapsulated powder have a higher modulus compared to the XY-plane. Again, the dry blend performed slightly better than the encapsulated material.

Again, the dry blend performed slightly better than the encapsulated material. The porosity found in the encapsulated material will have a negative impact on the compression modulus and can explain the observed difference.4000a P M3500/ s u u d o m n o i s s e r p m o Cl30002500200015001000Encapsulated PA613Dry blended PA613Neat PA613 [Kummert]5000X dryX condY dryY condZ dryZ condFigure 17: Compression modulus in dry and conditioned state of PA613 with dry blended and encapsulated ceramic spheres10080PA613 EncapsulatedC °/ A T D H6040PA613 dry blendNeat PA61320PA613 20 vol% glass beads0X dryX condY dryY condZ dryZ condFigure 18: Heat deflection temperatures at 1,8 MPa load of neat and filled PA613 powdersHeat deflection temperature of neat PA613 could be increased by 20 °C with the addition of ceramic spheres to neat PA613.

Both the dry blend and the encapsulated material exhibited the same heat resistance. While the heat resistance of glass a bead filled PA613 was reduced after conditioning the specimens, the ceramic filled PA613 showed isotropic heat resistance properties198that were not affected by conditioning. This indicated that the filler matrix adhesion of glass filled PA613 was impacted more severely by conditioning than that of ceramic filled PA613.Summary and OutlookWithin the scope of this work, two filled PA613 powders for laser sintering were investigated. Both powders consist of PA613 filled with around 35 wt% of ceramic spheres. One powder was produced by dry blending the ceramic spheres with commercially available VESTOSINT 3D 8754HT1 PA613 powder.

Both powders consist of PA613 filled with around 35 wt% of ceramic spheres. One powder was produced by dry blending the ceramic spheres with commercially available VESTOSINT 3D 8754HT1 PA613 powder. For the production of the other powder, the ceramic spheres were added during the dissolution precipitation process at Evonik so that the ceramic spheres were coated with AP613 powder and were therefore fully encapsulated.Analyzing the processability of both powders revealed, that encapsulating the ceramic spheres improved the powder flowability and provided a better dispersion of the filler particles within the polymer matrix. It was found that no segregation of polymer powder and filler occurred in the feed tanks of an EOS P3 system for the dry blend. In this regard, it is not necessary to encapsulate the filler particles. However, segregation effects could still occur in the powder recycling process during sieving.

In this regard, it is not necessary to encapsulate the filler particles. However, segregation effects could still occur in the powder recycling process during sieving. The encapsulated powder exhibited different thermal properties compared to the dry blend. A secondary melting peak before the main melting peak was identified as the cause for severe caking of the powder bed, even though the encapsulated material was processed at 18.5 °C lower powder bed temperature than the dry blend. With adapted processing parameters during the dissolution precipitation, it was possible to produce a new variant of the encapsulated powder that has almost the same thermal properties compared to the dry blend and will be investigated in further detail in the future.The resulting parts built from PA613 with encapsulated and dry blended ceramic filler have a white appearance with good resolution and surface finish.

Compared to neat PA613, both filled powders exhibited a higher tensile and compression modulus and an increased heat deflection temperature. Tensile strength and elongation at break were reduced for the filled powders compared to neat PA613. Some mechanical properties of PA613 with encapsulated ceramic filler were slightly worse than that of the dry blend. Especially in Z-direction the dry blend exhibited superior tensile strength and elongation at break. It is believed that this difference is caused by the higher porosity within the parts built with encapsulated powder. The porosity is most likely the result of the much lower powder bed temperature used to process the encapsulated material and might be improved with the new variant on encapsulated powder that does not show the secondary melting peak. It was found that the encapsulated material exhibits a more linear stress strain behavior than the dry blend. The same stress caused less strain.

It was found that the encapsulated material exhibits a more linear stress strain behavior than the dry blend. The same stress caused less strain. It is assumed that this improvement is due to a better filler matrix adhesion and/or a better filler dispersion of the encapsulated material. The results for this investigation with PA613 should be easily transferrable to PA12, since this polymer is also produced by a dissolution precipitation process at Evonik.199References[1] R. D. Goodridge, C. J. Tuck, and R. Hague, “Laser sintering of polyamides and other polymers,” Progress in Materials Science, vol. 57, no. 2, pp. 229–267, 2012, doi: 10.1016/j.pmatsci.2011.04.001.[2] J. Breuninger, R. Becker, A. Wolf, S. Rommel, and A. Verl, Generative Fertigung mitKunststoffen. Berlin, Heidelberg: Springer Berlin Heidelberg, 2013.[3] M. Schmid, Lasersintern (LS) mit Kunststoffen. München: Carl Hanser Verlag GmbH & Co.KG, 2022.[4] T. Wohlers, R. I. Campbell, O. Diegel, R. Huff, and J.

mospheric Control during Direct Selective Laser Sintering of Stainless Steel 314S powder.C. Hauser, T.H.C. Childs, K.W. Dalgarno, R.B. Eane. School of Mechanical. Engineering, The University of Leeds, UK.AbstractStainless steel grade 314S powders have been Selective Laser Sintered (SLS) in three different argon/air (oxygen) atmospheric mixtures. The amount of oxygen present during the heating, melting and fusing of the metal powder strongly limits the range of laser powers and scanning speeds for successful processing. As oxygen levels diminish, powder oxidation reduces. This reduces absorption of laser energy as well as balling and other detrimental surface phenomena. This paper reports the conditions for creating sintered layers and observations of part quality variation within these conditions.

This paper reports the conditions for creating sintered layers and observations of part quality variation within these conditions. Sintered microstructure observations are also helpful in determining thermal history changes.KeywordsAtmosphere, Oxidation, Surface Tension, Metallography.IntroductionThe SLS processing route of heating, melting and fusing metal powders over extremely short periods of time, in a back and forth (cyclic) manner is a unique manufacturing strategy. During metal. powder processing, oxygen can be ••·present in the surrounding atmosphere, contained within the porosity of the powder bed and in the form ofa passive layer of oxide on the surfaces of particles prior to sintering, the latterbeing common due to the large surface area associated with a powder mass. The presence of oxygen and variations in the temperature-time history during sintering can fuel surface oxide growth.

The presence of oxygen and variations in the temperature-time history during sintering can fuel surface oxide growth. This has led to a succession of reported detrimental effects during the SLS of a number of metallic powders [1,2]. However, surface oxide build up is not always a disadvantage. Other laser manufacturing routes (e.g. laser cutting) report considerable increases in the absorption of CO2 laser radiation on oxidised metal samples where normal conditions (no oxide growth) would strongly reflect 10.6Jlm radiation [3] ..•• Another concern is the liquid metal surface tension which influences the wetting angle between the solid and the liquid phases that can disruptbonding between rastered lines and individual layers [4].

Reports indicate that certain scanning and atmospheric conditions allow surface tension phenomena to dominate causing the melt pool to solidify into a series of balls in the wake of the travelling laser beam [2,5].This paper is for us a first step in understanding the process of metal powder melting, flow and solidification during SLS. The research asa Whole. is intended to clarify the affects and requirements of the SLS atmosphere. Particular attention hasfocusedonthe significance.of oxygen during the formation and· growth of the metal powder melt as the laser·· scans the powder bed surface over a range of laser powers and scanning speeds.265MaterialThe material under investigation was a standard grade gas atomised stainless steel powdercl~ssifi~das 314S HC'an&ustenitic steel with a ~ighcarboncon.tent (referred to as 314S}.·'l:'hematerialcomposition.is/giveninTable 1.

Thepowderwas sieved into fOlll" batches having particle size distributionsof-300+1501Jl1l,-150+75~m, -75+38~m and -38~m. Images of sectioned and polished samples of the pre-sintered powder were observed through a Scanning Electron Microscope (SEM) to help understand the re-solidification process after laser sintering.

Figure 1 shows two typical SEM images of the -150+751Jl1l powder batch: (a) shows the general powder shape while (b) indicates a mean grain size in a particle of ~5.01Jl1l.Bal.20.025.00.4Table 1: Stainless steel 314S HC material composition.Figure 1: (a) Typical size and shape of314S -150 +751Jl1l particles and (b) Grain size and shape within a single particle.SLS Machine and Environmental Control Equipment.The Leeds high power SLS.machine includes a 250W continuous waveCOz laserwith a spot diameter adjustable between 1.0mm and 2.0mm at the focal length.Galvanom~ter controlled mirrors direct the laser beam onto a 70.0mm diameter build area which is housed within a 0.03m3 (L=460mm, H=260mm, D=250mm) process cham.ber (Figure 2a) capable of sustaining a variety of atmospheric conditions including an absolute pressure of 10mbar. Oxygen dilution within the chamlJerCanbeachieved by a combination of evacuation followed by a continuous·purge of argon.

Oxygen dilution within the chamlJerCanbeachieved by a combination of evacuation followed by a continuous·purge of argon. (bottled argon at 99.9% purity) or by targeting argon directly ont%ut of the powder bed. The targeting techniques (Figure 2b) were: Method 1- by purging the build area with argon throu.gh a shroud which surrounds the powder bed; Method 2 - by percolation of argon through the powder bed. Both methods were tested.266Powder Bedis...'"Bo""...'" B'''''Powd'" BodC"'.. Pttspex ShrowlFOO'.dB«Iof Stllinlessst«ll ball bearings.Stainles.sStecI"""'''''Po_ lledViewing windowIPowder bed cylinder~Powder roller feed(a)... Argon Purge, ... Expelled Argon(1 )(b)(2)Figure 2: (a) the process chamber, (b-l) gas shrouding and (b-2) gas percolation system.Experimental ProcedureThe four batches of 3l4S were sintered in the conditions in Table 2.

Single line scans varying in length from l5.0mm to 50.0mm were sintered in three different atmospheric gas mixtures of air, air/argon and argon while observing changes in the melt quantity and quality, oxidation and oxidation related phenomena. The results were displayed as a series of sintering charts. Single layer sintering in the same conditions as the previous experiments then followed. Observations were recorded including influences of scan spacing on the changes in oxidation behaviour, on bond quality between rastered lines, on repeatability and on thermally induced distortion. The results were superimposed onto the previous constructed charts.The air atmosphere was achieved by exposing the powder bed to laboratory.air during sintering. The air/argon atmosphereyvas created by both shrouding and percolating techniques in Figure 2b. A five minute pre-sinterpurgei?f argon wasappliedbeforesintering under a continual flow of argon.

The air/argon atmosphereyvas created by both shrouding and percolating techniques in Figure 2b. A five minute pre-sinterpurgei?f argon wasappliedbeforesintering under a continual flow of argon. The argonatmospher~ (which still contained traces of oxygen) was achieved by evacuating the process chamber toa pressure of 50mbar followed by a pre-sinter purge of argon for 10 minutes. During sintering the gas inside the chamber was held at 30mbar above atmospheric pressure. The quantities of oxygen within the thre.e atmospheres a.pd the flow rates of the gasses were not measured. However, the threetypesofatmospher~were sufficiently diverse to allow judgement on the effects of oxyg~n depletion duringSLS.The powder was in its as received state. No powderpre-heating",,~scarri~doutprior to sintering. However, only powder from containers with unbrokel1sealsi""asvsed:prolonged atmospheric exposure caused powder.

No powderpre-heating",,~scarri~doutprior to sintering. However, only powder from containers with unbrokel1sealsi""asvsed:prolonged atmospheric exposure caused powder. agglomeration from moisture absorption in the -38J.! batch and increased the chances offurther oxidation ofall powders during sintering.Exp. No.1 2 3Atmospheric ConditionAir Air/Ar20n ArgonBeam Diameter (mm) 1.1 1.1 1.1Laser Power (W)Scanning Speed (mm/s)0 - 185 in steps of ~10 0 - 185insteps of ~10 0 -185 in steps of ~10 Table 2: Experimental conditions.O~ 50 in st~ps of~2 0 ~ 50inst~sOf'-2 0 - 50 in steps of ~2Scan Spacing (Fraction of beam diameter)0.25,0.5,0.75 0.25,0.5,0.75 0.25,0.5,0.75267Sintering Maps and Observational ResnltsExp. 1: Air atmosphere sintering.Conditions for sinterillg single lines and single layers within an air atmosphere were found to be limited.

1: Air atmosphere sintering.Conditions for sinterillg single lines and single layers within an air atmosphere were found to be limited. Figure 3 maps the observed stages of oxidation b~haviour, particle bonding .and melt pool growth and quality during single line scanning and Figure 4 shows typical examples of single layerscanning using Figure 3 as areference guide. During sintering temper colorations following classical oxidation theory [6] could be observed on the surfaces of individual 314S particles as the powderbed was heated by the laser beam. As the net energy density increased (power/speed), small droplets· (balls) of liquid metal began to form covered by an opaque surface oxide scale and seated within a trail of oxidised powder. The phenomenon of melt pool balling and breakage was widespread during air sintering because liquid surface tension forces were allowed to dominate due to large melt volumes and the influence of the surface scale.

If the scanning speed remained low «8.0mm/s) and the laser power was high (dependant on speed) surface tension forces became less dominant and the melt pool began to flow more freely. This created a continuous melt pool that solidified into a tubular shape, though a single ball did exist at the start (this can be seen in Figure 4). However, such high temperature conditions dramatically increased the melt volume and a slag believed to be rich in chromium (Cr203) was observed on the surface.Single layer scanning was carried out over the range of conditions where single linescanning was successful (white zone of Figure 3). However, it was found that single layer300 + 150J.Ull150 + 75Jlrn~~..

However, it was found that single layer300 + 150J.Ull150 + 75Jlrn~~.. J 50L e melt VolumeContinuous melting and bonding exhibiting high surface tension phenomena Full melting with breakage'sSunning Speed (tom/II)Scanning Speed (mmls)1111 Balling75 + 38J.Ull i 'Temper Coloration's and partial melting (Bonding).150Temper Coloration's (No Bonding).~ i l00Il~'~~-i··~~i~' .. i'"'50·No MarkingBounded area where single layer sintering is possible.•10Scanning Speed (mm/s)s.:D"'UJg S"",d (mmi.)Dependant on scan spacing.Figure 3: Sintering maps for 314S processed within an air atmosphere.268ISO + 7SIJl1lScan Spacing 0.7S'Scan Spacing 0.5'Scan Spacing 0.2S'Figure 4: Examples of air atrn.osphere sintering.scanning could only be achieved at very high energy densities. Lower energy densities produced a continuous First Line Scan (FLS) followed by balling of all other rastered lines (See Figure 4).

Lower energy densities produced a continuous First Line Scan (FLS) followed by balling of all other rastered lines (See Figure 4). Experimentation also showed that if the scan spacing was large (>0.6 x width of single line) then bonding was limited and if the scan spacing was low «0.4 x width of single line) balling occurred.Exp. 2: Air - Argon atmosphere sintering.Figures 5 and 6 show the process map for single line air-argon sintering and typical examples of single layers sintered in an air-argon atmosphere respectively. A similar trend to thatof air atmosphere processing is seen. However single layer scanning was more successful with improved bonding between the rastered lines over·a larger range.of scanning conditions, including more flexibility in the choice of scan spacing. The ball at the start of the scan length had also dramatically reduced. Warping of the.singlelayers was also evident, showing that large bulky layers are. affected by thermal distortions.

The ball at the start of the scan length had also dramatically reduced. Warping of the.singlelayers was also evident, showing that large bulky layers are. affected by thermal distortions. All results were obtained from the gas shrouding equipment. •When using the gas percolator powder was blown from the bed.Exp. 3: Argon atmospheresintering.Figures 7 and 8 show examples of argon atmospheresinteringand the process map for single linesintering within an argon atmosphere respectively. Under these conditions no oxide growth occurred resulting in free flowing smaller melt volumes. creating flatter (O.4mm to 1.3mm thick) and stronger bonded single layers. llowever, because of the lower meltvolume, porosity still existed, effectively lowering thesintereddensity by 40 to 50%.

creating flatter (O.4mm to 1.3mm thick) and stronger bonded single layers. llowever, because of the lower meltvolume, porosity still existed, effectively lowering thesintereddensity by 40 to 50%. Even extremely high energy densities (185W at induce> full melting as previously lImn/s) could not experienced in air and air - argon sintering.Metallographic InspectionFigure 9 shows typical results of sectioned,. polish~dand etched single sinteredlines showing how the microstructure and cros~sectional s~ap~chat1~ewithchanges in sintering solidification modes were observed. Firstly, a cellular atmosphere. mJcrostructure>Jound in both air (full) and air/argon (upper halt) sintering indicating re solidification from temperatures sufficient to induce full melting. Secondly, an equiaxedTwo different269300 + 150!!ffi200ISOLame melt Volume~ i 100 ..

temperafures within the solidus/liquidus or 'mushy' zone of the phase diagram where solid and liquid co-exist. The shape of the bqundary is a good indication of the changing dominance of the acting surface tension.force.. Figure9.also.indicates the height of sinteredsingle lines (also reflective. of single la.yers) relative to the powder bed surface. Sinking of the sinteredpartwas observed as the melt volume increased. This phenomenon can illfluenceconditions during powder recoating.Final Discussion and ConclusionsThe presence of oxygen within the sintering atmosphere and powder bed allows surface oxides and slags to form as the powder is heated and melted by the scanning laser beam. Observing the growth and effects of these scales in differing atmospheric conditions has shown that elimination of all oxygen is required in order to reduce the melt volume allowing surface tension forces to become less dominant.

Under these conditions liquid metal can flow and wet other solid powder particles and previously sintered areas creating strong bonds for successful single layer sintering. Average sintered densities ranged from 50 - 60 % of theoretical density. However higher densities of ~70% were attainable when using a smaller particle size(-38Ilm batch) and high scanning energy densities. The existence of surface oxides and slags increased the melt pool volume dramatically (5mm thick single layers were produced). Possible reasons for this include: (1) an increased coupling efficiency between the laser and the surface oxide layer, (2) the exothermic nature of some oxides and (3) the oxide acting as an insulating jacket.

The large melt volumes solidified to full density at the expense of a dominant surface tension211TemperatureLaser Power (P) at speed (u).Level ofPowder BedIlmm IResults obtained from -150 + 75,.un batch at l20W and 2mm/s.- _--~,--,"Air Processing -_Air- Argon Processing -_~_.- --_ Boundary between cellular and__..-_. i dendritic ~icrostructure. (Represents Liquidus Linei[. i : ion phase dIagram)"',Location of Bobble( Jd= f :)Argon Processing.......-Laser...... ________________JFigure 9: Changes in microstructure and shape of single sintered lines.force which lowers wetting causing sharp reductions in the range of available scanning conditions for single layer sintering. Improved wetting and bonding was achievable in the air/argon atmosphere. However metallographic inspection has shown that bonding only occurred in areas between adjacent scans where surface tension phenomena were relaxed i.e.

However metallographic inspection has shown that bonding only occurred in areas between adjacent scans where surface tension phenomena were relaxed i.e. in regions where solid fragments and liquid co-exist.AcknowledgementsThe work reP9rted in this paper was supported by the UK Engineering and Physical (EPSRC), through an IMI grant and industrial collaborators. TheScience Research Council authors would like to thank the EPSRC and industrial partners for there support.REFERENCES.[1 ].[2].[3]. [4].[5].[6].U. Lakshminarayan et aI, Advances in Manufacturing Metal Objects by Selective Laser Sintering (SLS™), Advances in Powder Metallurgy and Particulate Materials, Volume 4, 1996. W. T. Carter et aI, Direct Laser Sintering of Metals, Solid Freeform Fabrication Symposium, The University of Texas at Austin, 1993 [pp 51- 59]. G. Chryssolouris, Laser Machining: Theory and Practise, Springer-Varlag, 1991. J. A.

TOMATED 4 AXIS ADAYfIVE SCANNING WITH THE DIGIBOTICS LASER DIGITIZERINTRODUCTIONThe DIGIBOT 3D Laser Digitizer is a high performance 3D input device which combines laser ranging technology, personal computing, and Microsoft Windows in an attractive desktop package.the DIGIBOT provides a simple, accurate, and quick way to copy or inspect complex, sculpted surfaces. The DIGIBOT provides an effective solution for many industrial and academic problems involving 3D inspection, replication, analysis, and visualization/animation.

The DIGIBOT provides an effective solution for many industrial and academic problems involving 3D inspection, replication, analysis, and visualization/animation. By measuring design, sequential points and producing a standard list of x/y/z coordinates, the DIGIBOT interfaces to any CAD/CAM/CAE imaging or animation software that reads 3D points, contours, or triangular facets.With its full four-axis scanning capabilities,UNIQUE BENEFITS OF THE DIGmOT SYSTEMThe following unique benefits are provided by the DIGIBOT technology:DIGIBOT's ranging system measures individual surface points without using imaging optics or sophisticated detector-array processing complex, non-linear techniques.The ranging system provides high resolution and accuracy while maintaining a small, fixed triangulation angle across a large work volume...

Unlike camera-based ranging devices, the Digibot Ranging System has a very long ranging length with a small stand-off and does not suffer from focal problems such as lens aberrations and depth-of-field issues. Systems can be mechanically configured to provide up to two rotations and two translations between the object and the ranging system. possibleAdvancedpositioningcapabilities makethecomplete,one-scanmeasurement of complex objects.Adaptive 4 axis scanning procedures minimize shadowing effects, provide optimal spacing with no measurement orientations, and produce homogenous point measurement redundancy. Semi-adaptive, systematic, and interactive scanning procedures provide efficient alternatives for less sculpted, more regular surfaces. A complete library of C-based system functions can be used to develop custom scanning procedures.The system is a peripheral device which interfaces to an IBM compatible PC through a controller board located inside the PC.

The system is self-orienting and maintenance free. The Digibot calibration/registration procedure is quick and easy to the procedure calibration/registration processes used for camera-based systems. less150DIGIBOT systems are easy to set up and move because of their small size and low weight.Digibot' s windows-based software provides a conventional, intuitive interface to the scanning procedures and a variety of data processing utilities, including contour editing, surface editing, polygonal surface generation, adaptive surface filtering, 3D viewing, and data formatting. DIGITIZING DESCRIPrIONDigitizing an object begins by first mounting the object on the platter in front of the laser scanning system. The operator then uses the Windows interface to select from a variety of scanning procedures, each of which will be more or less appropriate for a given object. Once selected, the operator instructs the system to execute the scanning procedure.

Once selected, the operator instructs the system to execute the scanning procedure. Typical scanning procedures, systematic or adaptive, provide sampling rates on the order of 10,000-100,000 points per hour. A course sampling grid can be executed in minutes while high resolution sampling of complex surfaces can take hours.DATA (GEOMETRY AND TOPOLOGY)including the control software resident on the customer's host computer is essentially an automatic data acquisition system. Typically, the next step after acquiring data with the digitizer is to manipulate the data for some purpose such as graphical display or numerical analysis. At this time the user must understand and appreciate the 'art and science' of using a set of discrete surface points to mathematically represent a surface. In essence, the 3D coordinate values of individual surface points only provide information about the geometry of a surface. A complete surface representation also requires topology (i.e.

In essence, the 3D coordinate values of individual surface points only provide information about the geometry of a surface. A complete surface representation also requires topology (i.e. information about how each point is connected to its neighboring points). The selection of one scanning procedure over another will typically have a great influence on the amount of explicit topology information that can be acquired during the digitizing process. Some scanning procedures can provide complete topology while others provide only partial or no topology. Any data that is lacking in topology information must typically be subjected to some form of mathematical topology generator. A procedure of this type must In general, the user must be attempt to establish topology from the available geometry.

A procedure of this type must In general, the user must be attempt to establish topology from the available geometry. aware that the selection of one scanning procedure over another also involves a conscious decision to use one data processing procedure over another.The DIGIBOT device,(IBM PC),ADAPTIVE SCANNING AND FILTERINGWhen measuring an analog subject, it is generally a good idea to perform a high resolution sampling procedure so that high frequency components are adequately resolved. This principle is appropriately applied when digitizing a 3-dimensional object. The representation of a sharp edge will typically require a dense grid of closely spaced surface points. Although essential, high resolution scanning procedures also present a problem. Slowly curving surfaces, which can be represented with relatively151few widely spaced surface points, tend to be over-sampled producing excessive amounts of unnecessary data.

Slowly curving surfaces, which can be represented with relatively151few widely spaced surface points, tend to be over-sampled producing excessive amounts of unnecessary data. The solution to this problem is provided by introducing some form of filtering mechanism which uses the local rate of surface curvature to adaptively adjust the spacing between neighboring surface points. With the DIGIBOT system, a filtering procedure of this type can be introduced in two ways. When digitizing, an adaptive scanning procequre can be executed so that the spacing between consecutive measurements is decreased for increasing rates of surface curvature. After digitizing, the acquired data can be subjected to a flexible filtering procedure which removes unnecessary points in over-sampled surface regions. Generally, the second approach is the most attractive and flexible (Le.

Generally, the second approach is the most attractive and flexible (Le. it is the best to over-sample the object and then use a data filter to selectively remove a desired set of surface points).POINT MEASUREMENT BY TRIANGULATIONIndividual range measurements are obtained using an active triangulation method. The ranging system is composed of a laser source and a translating photo-direction laser source and the photo-direction detector are positioned at two detector. The comers of a right triangle formed by the orthogonal intersection of the laser line and the base line. At the third comer of the triangular system the laser beam illuminates a small surface producing a diffuse point-source of light. The photo-detector is translated until the illuminated surface point enters the detector's field of view.

The photo-detector is translated until the illuminated surface point enters the detector's field of view. At this time the system can accurately determine the distance and the angle theta from which the range can be computed.Active triangulation requires that the surface be a relatively diffuse reflector and that the angle of incidence between the laser line and the surface normal be within limits. Polished surfaces cannot be measured because, in most orientations, no light will be reflected back toward the photo-detector. Increasing incident angles will tend to illuminate an increasingly larger surface area while decreasing the amount of light reflected back toward the photo-sensor.There is a limit to the effectiveness of a triangulating range system. As surface concavity increases, the photo-detector's view of points located within concave regions becomes increasingly obstructed, producing increasing limited laser/photo-sensor orientations.

As surface concavity increases, the photo-detector's view of points located within concave regions becomes increasingly obstructed, producing increasing limited laser/photo-sensor orientations. tend to decrease measurement resolution and accuracy. Typical applications will use a 30 degree probing angle.Reduced probing angles provide improved probability butRELATIVE MOTIONS (DOFs) BETWEEN LASER BEAM AND OBJECTAssuming that the ranging device has a long ranging length and does not need to move toward or away from an object, there remains four forms of relative motion between an object and a laser beam needed to illuminate a point on any exposed surface of a 3-dimensional object. More specifically, a 3D digitizer needs two translational and two rotational degrees-of-freedom (DOFs) between the ranging system and object to provide full 3D sampling capabilities. DIGIBOT offers two translational DOFs and one rotational DOF. This is also known as four axis scanning.

DIGIBOT offers two translational DOFs and one rotational DOF. This is also known as four axis scanning. The laser can be152translated vertically and horizontally while the object can be rotated about a vertical axis. By not providing the fourth DOF (i.e. rotation of the object about a horizontal axis), DIGIBOT is not capable of measuring points along vertical facing surfaces. For most applications, the practicality of using a 4-DOF digitizer is questionable due to mechanical, operational, and data processing difficulties. The fourth DOF can, however, be partially realized by digitizing the same object multiple times in different orientations. This approach requires that the multiple data sets be merged to form a single surface mesh.By using different combinations of its three DOFs, DIGIBOT is capable ofperforming three types of 3D scanning procedures:1. 2. 3. Each of these three scanning options is useful for different types of objects.

2. 3. Each of these three scanning options is useful for different types of objects. Flat objects are best digitized using two translational DOFs while convex objects are best digitized using both translation and rotation. Three DOF scanning offers a very attractive method for digitizing complex objects with non-axial geometry, multiple contours, and concavities. The following three discussions provide a more detailed description of when and how the different scanning options can be used to digitize different types of objects. SCANNING WITH TWO TRANSLATIONAL DOFstwo DOF scan using two translational DOFs, two DOF scan using translational DOF and rotational DOF, and three nOF scan using two translational nOFs and one rotational nOF.Probably the simplest scanning procedure that can be performed with nIGIBOT employs two translations of the ranging system. As a result, this type of scanning option should be used to measure relatively flat surfaces.

As a result, this type of scanning option should be used to measure relatively flat surfaces. Once the user has decided which combination of motions are appropriate for a given application, a decision must be made to perform either a systematic scanning procedure or an adaptive scanning procedure.The surface mesh generated from adaptive data is very attractive because it produces a homogeneous, evenly spaced set of points on all exposed regions of a surface. An adaptive mesh can also be optimized to provide more points in high in regions where high sample density is frequency regions and fewer points unnecessary. Though the data produced by adaptive scanning procedures is attractive, in practice it is generally more practical for simple-shaped objects to employ fully systematic procedures with only two DOFs.

Though the data produced by adaptive scanning procedures is attractive, in practice it is generally more practical for simple-shaped objects to employ fully systematic procedures with only two DOFs. The reasoning for this is two fold:1.the grid topology of systematic data is provided implicitly while the mesh topology of adaptive data must, by some method, be explicitly specified and recorded (this can be difficult), and2. systematic data from a high density grid can be adaptively filtered to produce an adaptive surface mesh that is very similar in quality to that produced by adaptive scanning procedures.Though adaptive scanning methods are used for most applications, systematic scanning procedures should be used whenever possible because they are easy to perform and they provide structured data that can be flexibly filtered to provide an153optimal data sub-set.

Another advantage to using high frequency systematic sampling procedures is realized when the user decides that more data is needed. It is generally easier to refilter a set of data than to re digitize an object.SCANNING WITH ONE TRANSLATIONAL DOF AND ONE ROTATIONAL DOFA very popular form of 2-DOF scanning utilizes both rotation and translation. By first positioning the laser line to pass orthogonal through the center of rotation, a set of radial measurements can be obtained by rotating the object and vertically translating the ranging system. This scanning configuration is useful for convex objects because acceptable incident angles are possible only at locations where the surface normal does not depart significantly from the radial direction. When choosing between systematic introduced in the versus adaptive scanning procedures, previous discussion describing 2-translation scanning, can again be used.

When choosing between systematic introduced in the versus adaptive scanning procedures, previous discussion describing 2-translation scanning, can again be used. A systematic scanning procedure will produce a data-grid (with complete topology and filtering capabilities) by incrementally rotating the object and incrementally translating the ranging system.

Some subtle properties of rotation contour sampling are:the very same reasoning,points with a surface tangent that passes through, or near, the center of rotation cannot be measured, negative radiuses can be measured because the active range of the ranging system extends past the center of rotation, closer spacing between measurements occurs along contours that are closer to the center of rotation, off-center objects are sampled during two portions of a complete rotation causing partial loss of topology and partial loss of the contour on two sides where large incident angles make measurement impossible, and • most concave surfaces cannot, or can only partially, be measured.Two DOF scanning procedures are very attractive because they are easy to perform and they produce data that is easy to work with. Unfortunately, the type of object that can be digitized using 2-DOF methods is limited.

Unfortunately, the type of object that can be digitized using 2-DOF methods is limited. With its third DOF, DIGIBOT is capable of digitizing much more complex objects, including off-center contours, multiple contours, and concavities. However, as a general rule, 2-DOF scanning procedures should be used whenever possible to take advantage of gridded data and to avoid the difficulties associated with the processing and compatibility of 3-DOF data. reasoning behind this rule will become clearer after reading the followingdescription of 3-DOF scanning.154SCANNING WITH ALL THREE DOFs (4 AXIS SCANNING)The work volume generated using all three OOFs (Le. rotation of the object and both vertical and horizontal translation of the laser line ) will be cylindrical in shape. With three OOFs, OIGIBOT provides a work volume with an 18" diameter and a 18" height.

With three OOFs, OIGIBOT provides a work volume with an 18" diameter and a 18" height. The advantage of using a third DOF is two fold:1.the surface normal at any given point is limited only by the requirement that it not depart significantly from the horizontal, and2. each individual point can be measured from a variety of different perspectives the laser beam can illuminate a single surface point from a variety of (Le. different directions).Complex contours can be better sampled along a horizontal plane by both rotating the object and translating the ranging system. Deep concavities and multiple contours can only be measured using both rotation and translation while, for complex contours of this type, rotations by itself will yield comparatively poor results.An efficient 3-DOF procedure for sampling contours along a horizontal cross section requires adaptive feedback control.

The direction and rate of curvature along a contour are computed using previously acquired points to help predict the location and surface normal of a target point (Le. the next point to be measured). In this way, the rotated and the ranging system is translated to provide an optimal object measurement perspective (Le., the laser as normal as possible to the surface) for each For convex consecutive target point as the system adaptively tracks a contour. surfaces, the optimal measurement perspective is obtained with a zero incident angle between the laser beam and the surface normal. For concave surfaces, optimal measurement perspectives are obtained when the laser line is closely pivoted about the outer edges of the concavity. Orientations of this type will typically produce non-zero incident angles. A very useful result of adaptively moving along a contour can be seen by measuring consecutive points. The system effectively acquires complete topology information (Le.

A very useful result of adaptively moving along a contour can be seen by measuring consecutive points. The system effectively acquires complete topology information (Le. the points are connected in the same order in which they are acquired) and, when multiple contours exist, each contour is scanned separately (Le. once started, the systems attempts to complete a contour by pivoting around any other obstructing contours.)isThough two DOFs are used to scan adaptively within a horizontal plane, the third DOF is typically used to systematically position the ranging system at vertically displaced planes. For this reason, this type of 3-DOF scanning could be referred to as It is important to recognize that adaptive scanning along planar cross 'semi-adaptive. I sections provides complete contour topology in the horizontal direction. This is a great advantage but, by blindly moving to vertically displaced planes, semi-adaptive scanning will provide no vertical (Le.

This is a great advantage but, by blindly moving to vertically displaced planes, semi-adaptive scanning will provide no vertical (Le. no information is provided to establish how points in one planar contour are connected to points in a corresponding, vertical-displaced planar contour). As a result, a procedure for vertical topology generation must be employed before a valid surface mesh can be obtained from semi-adaptive data.topology between neighboring cross-sections155A filtering mechanism can be incorporated into the procedure for semi-adaptive scanning to provide a flexible way to discriminate between rapid and slow rates of surface curvature. By using the local rate of surface curvature to adjust the spacing between consecutive measurements, a relatively small number of measurements are needed to define a contour containing sharp corners and slow curving sides.

Without a filtering mechanism, a contour of this type must be sampled frequently to guarantee Such a procedure will require that the sharp corners will be adequately resolved. significantly more time and will, in this case, produce an over-abundant amount of data along slow curving surfaces.There is, unfortunately, a competing disadvantage to using an adaptive filtering mechanism along planar cross-sections when performing a semi-adaptive scanning procedure. By systematically stepping to consecutive planes, no filtering mechanism is applied in the vertical direction producing a non-homogenous surface mesh composed of long, skinny polygons. This problem can be resolved by employing a vertical filter either during or after the scanning procedure. When scanning, a fully adaptive procedure is fast and concise but it is difficult to implement and the resulting data must explicitly specify topology.

When scanning, a fully adaptive procedure is fast and concise but it is difficult to implement and the resulting data must explicitly specify topology. After scanning, a vertical data filter must first be applied to produce a homogenous surface mesh which will also require explicit topology information. As a consequence, unless time is a factor, or for some other specific reason, it is generally more practical to execute a semi-adaptive scanning procedure with equal sample spacing in both the vertical direction and along horizontal contours. The data produced from such a procedure can be subjected to a vertical topology generator, and then filtered to extract the desired data subset.SUMMARYThe Digibot 3D Laser Digitizer is a high performance 3D input device which combines laser ranging technology, personal computing, and Microsoft Windows in a desktop package.

With its full four-axis scanning capabilities, the Digibot provides a simple, accurate, and quick way to copy or inspect complex, sculpted surfaces. The Digibot provides an effective solution for many industrial and academic problems involving 3D design, inspection, replication, analysis, and visualization/animation. By measuring sequential points and producing a standard list of x/y/z coordinates, the Digibot interfaces to any CAD/CAM/CAE imaging or animation software that reads 3D points, contours, or triangular facets.The DIGIBOT uses a unique triangulation technique that does not use complicated imaging optics or array detectors. The scan head resides outside of the scan region and does not move back and forth from the object to maintain focus of the laser beam spot.

The scan head resides outside of the scan region and does not move back and forth from the object to maintain focus of the laser beam spot. Consequently, Adaptive scanning procedures intelligently position the beam spot on the object's surface while effectively walking from one point this technique, unique to the For complex objects, DIGIBOT, can reach into deep concavities, undercuts, and between surfaces to produce a complete, homogeneous data mesh. Simple systematic scanning procedures produce grid data for single faced (Le., flat-faced) objects or convex, cylindrical shaped objects.scanning is simple and efficient.to the next.156Three dimensional modeling is becoming the technology of choice for the majority of designers and manufacturers. In order to be effective, many industries need to automate the acquisition of 3D geometry into the computer. The Digibot solves this need with industry leadership in technology, usability, and affordability.157.

However, there are key challenges regarding the feasibility of current image-guided monitoring methodologies in the DED process. First, high-resolution thermal images consist of millions of pixels captured by hundreds of frames lead to the curse of dimensionality in analysis. Second, the presence of various exogenous noise, ill-structured data, and significant cluster imbalance limit the capability of the current methodologies to perform real-time monitoring. The objective of this research is to advance the frontier of melt pool monitoring in DED process by designing an automated and unsupervised anomaly detection on high-dimensional thermal image data. Specifically, we develop a variational autoencoder to generate a low-dimensional representation of each input thermal image data. A Gaussian mixture model and K-Mean clustering are integrated with the generative model to split latent space into homogenous regions and detect anomalies.

A Gaussian mixture model and K-Mean clustering are integrated with the generative model to split latent space into homogenous regions and detect anomalies. Experimental results show that the proposed methodology is highly effective in detecting defective melt pools with accuracy up to 94.52% and a false alarm rate of less than 2.1%.IntroductionThe ability of metallic additive manufacturing (AM) to produce intricate geometry parts (e.g., lattice structures and internal channels) from hard-to-process materials (e.g., Ti-6Al-4V and 17- 4PH alloys) manifests their potential to revolutionize fabrication. This popularity is further driven by virtue of higher design flexibility, shorter development time, lower tooling cost, and less production waste. The manufacturing flexibility offered by AM is valuable in a variety of strategic applications ranging from aerospace to biomedical.

The manufacturing flexibility offered by AM is valuable in a variety of strategic applications ranging from aerospace to biomedical. For instance, using AM to make components for the Cessna Denali aircraft engine reduced the number of components from 855 to 12, and improved the fuel efficiency of the engine, as well as its power by over 10 percent [1]. However, the significant challenge of metal AM is the occurrence of various defects such as cracks, delamination, distortion, rough surface, lack of fusion, porosity, foreign inclusions, process instability (keyhole and balling), which in turn, deteriorate the builds strength, residual stress, hardness, and fatigue life [2-4]. As a result, the quality and reliability issues seen in many metals AM hamper further proliferation of this technology in mission-critical applications [5].Porosity is the prominent type of defect in AM.

The occurrence can be traced back to rapid solidification of AM parts, entrapped gas, incomplete powder melting, and lack of fusion [6]. The377thermal distribution of melt pool has received significant attention as a quality indicator since it provides underlying information about the complex physical and metallurgical processes, including heating, melting, Marangoni convection, evaporation, solidification, and morphology. In particular, the abnormal behavior of melt pool can be directly correlated to the formation of porosity, the most common type of defect in metallic AM.Current defect detection methodologies are (1) offline (X-ray computed tomography (XCT) and ultra-sonic scanning) or (2) lumped-mass formulations, which are time-consuming, costly, and less effective with large-size builds [7].

Advancement in in-situ sensing technology provides an unprecedented opportunity for on-the-fly melt pool monitoring in various AM processes such as powder bed fusion (PBF) and direct energy deposition (DED). Significantly, the DED process offers the capability to produce and repair high-value and large-size components using Laser Engineered Net Shaping (LENS) and Direct Metal Deposition (DMD) systems. However, they suffer from a lack of part consistency and quality.Images of the melt pool capture key thermal activities during the DED process, thus providing the most direct evidence of microstructural anomalies. For example, the small melt pool connecting to adjacent melt pools may cause porosity defects due to the presence of unmelted powder. Also, a large melt pool may develop remelting and overheating problems.

Also, a large melt pool may develop remelting and overheating problems. The current image-guided monitoring focuses on analyzing the morphological features (e.g., length, depth, area, and size) or thermal distribution (average and maximum temperature) of melt pool. Therefore, the detection capability significantly relies on the quality of handcrafted features used in unsupervised learning models such as K-means [8], Gaussian mixture model (GMM) [9], and spectral clustering [10]. Recently, deep neural networks (DNNs) have shown remarkable ability in image-guided learning and autonomous feature extraction [11-13]. DNNs are less dependent on domain knowledge and simultaneously learn morphological features and thermal distribution through backpropagation by utilizing various building blocks (e.g., fully connected, convolutional, and pooling layers). However, there are fundamental challenges to the-state-of-the-art monitoring methods using thermal image streams:1.

However, there are fundamental challenges to the-state-of-the-art monitoring methods using thermal image streams:1. High dimensionality and velocity: high-resolution images are comprised of thousands and millions of pixels. A standard video camera collects 24 frames per second, whereas a high- speed camera may acquire thousands of frames per second, which is a hurdle for real-time and image-guided monitoring.2. Data structure: The low signal-to-noise ratio due to the DED by products (e.g., spatters, splashes, and fumes) significantly challenges the meaningful extraction of information. Furthermore, ill-structured data (e.g., different melt pools varying in sizes) and significant cluster imbalance (i.e., low number of defective images compared to the entire data) introduce uncertainty into the analysis.

Hence, there is a dire need for robust methodologies that account for this uncertainty and realize real-time melt pool monitoring in DED.The present research investigation is aimed at filling these gaps by developing an integrated deep generative model and unsupervised learning to automatically encode and discriminate meaningful patterns of variation in morphology and thermal distribution of images with significant cluster imbalance. Specifically, we design a variational autoencoder (VAE) model consists of an encoder (i.e., a recognition network) that converts the input to a latent representation and a decoder378(i.e., a generative network) that transforms the internal latent representation to the outputs. The VAE forms an information bottleneck, compresses high-dimensional image data into a low- dimensional latent space, and forces the encoder to extract the most critical features of defects into a continuous and meaningful embedding.

Then, the GMM and K-mean clustering as unsupervised models are integrated with the variational information bottleneck to disentangle melt pool image data through useful discriminative features.We evaluate and validate the proposed methodology with a real-world case study on thermal images captured from an OPTOMEC LENS 750 machine equipped with 1-kW Nd:YAG laser, pyrometer, and in-chamber thermal camera, which is used to build single track Ti-6A1-4V thin walls in Mississippi State University. Also, XCT of specimens are utilized to compare the predicted locations of pores based on the proposed methodology with their actual locations. The in- situ sensing and process monitoring and control are the next vertical steps to mitigate scrap and rework rates and further ensure the economic viability of AM.

The in- situ sensing and process monitoring and control are the next vertical steps to mitigate scrap and rework rates and further ensure the economic viability of AM. This project will eventually lead to the minimization of porosity during manufacturing as well as the maximization of final part strength and fatigue resistance.The remainder of this paper is organized as follows: We first provide a literature review on the relevant methods of sensor-based flaw detection and melt pool characterization. Then the proposed deep generative model and unsupervised learning methodology are provided. Experimental design and materials along with experimental results on thin-wall structures are given in the next two sections.

Experimental design and materials along with experimental results on thin-wall structures are given in the next two sections. In the end, we conclude this paper by highlighting gaps in existing characterization methodologies for defect monitoring, then provide an overview of the proposed methodology.Research BackgroundSensor-based flaw detection:Although empirical process mapping approaches have been utilized to avoid porosity, the occurrence of arbitrary defects is prevalent due to the dynamic condition of the DED process. Such random defects demand expensive and time-consuming post-build characterization. The XCT scanning and ultrasonic signal are among the most common offline sensing methods for defect characterization. The ultrasonic analysis is primarily investigated to quantify the mechanical strength and porous structure and find internal defects such as cracks, voids, delamination, to name a few [14].

The ultrasonic analysis is primarily investigated to quantify the mechanical strength and porous structure and find internal defects such as cracks, voids, delamination, to name a few [14]. However, it is sensitive to the selective layers of samples and ineffectual to capture pores between layers [15]. XCT has been extensively utilized not only to characterize the porosity level of build but also to link the variation in process parameters to the quality of final build [16, 17]. The limitations of the offline characterization methods are associated with cost and required time for analysis and the time restraint to leverage the unique layer-by-layer building mechanism of AM for in-process correction/control of defects before the next layer is built.In-situ detection of defects using sensor signatures suggests the promise of reduced inspection cost and increased confidence in part quality.

Five main types of signals are usually collected during the fabrication process; namely, visible lights and ultraviolet (UV) emitted from the plasma, IR emitted from the melt pool, ultrasonic waves from solid structure, audible sounds from air, and electric signals from plasma [18]. These signals can be captured with the photodiode, high-speed charge-coupled device (CCD) camera, complementary metal oxide semiconductor (CMOS) camera, and acoustic transducer [19, 20].379Acoustic signals are collected from the plasma that is generated at the powder bed surface. Using underheating or overheating of metal powder, the surface temperature of parts varies and causes a change in plasma density. Fluctuation of atmospheric pressure along with variation of plasma density impact the acoustic intensity and can be exploited as a process signature [21].

Fluctuation of atmospheric pressure along with variation of plasma density impact the acoustic intensity and can be exploited as a process signature [21]. Although acoustic signals can also provide a high temporal resolution for tracking the location of defects more accurately, the background noise is noticeable for many AM processes, particularly for the laser metal AM processes that leverage inert gas in their chambers [22]. Optical-based techniques are the most widely adopted in-process monitoring methods in the DED community. Digital cameras, high-speed cameras, and infrared (IR) thermal cameras are commonly used to capture optical signals. In some cases, photodiodes and pyrometers can provide supplementary information [23]. Note that CCD or CMOS cameras fitted with near infrared (NIR) filters or infrared pyrometers are implemented to measure the shape, size, or temperature of the melt pool [24].

Note that CCD or CMOS cameras fitted with near infrared (NIR) filters or infrared pyrometers are implemented to measure the shape, size, or temperature of the melt pool [24]. Photodiodes are integrated with the tube or nozzle to quantify the powder flow rates [25]. A camera is employed to ascertain the clad height in DED [26]. In addition, optical emission spectroscopy has been applied to monitor the optical emission above the melt pool that is created by a high-energy vapor plume [27].Melt pool monitoring and characterization:Melt pool images are the most informative process signatures for in-situ monitoring, as the shape and temperature of melt pools determine the occurrence of defects and dimensional accuracy. Therefore, various researchers have been designed statistical and machine learning methods to leverage sensing data from the melt pool for defect detection [28]. Khanzadeh et al. [29] investigated the effect of the heated zone on the emergence of flaws in DED.

Khanzadeh et al. [29] investigated the effect of the heated zone on the emergence of flaws in DED. They demonstrated that by monitoring the features of melt pool images obtained from functional principal component analysis (FPCA) of dual-wavelength imaging pyrometer data, the onset of lack-of-fusion defects could be predicted through machine learning. Grasso et al. [30] integrated the plume images from an IR camera with a sampling frequency of 50 Hz and a spatial resolution of 320 × 240 pixels. The grayscale normalization was performed, and the regions of interest were extracted by image thresholding and segmentation. As such only the relevant portion of the image was used to effectively reduce computational cost and processing time of monitoring.Supervised ML methods, including multilayer perceptron (MLP) and support vector machine (SVM), were applied to differentiate the distinctive thermal signatures of melt pools of overhang sections from bulk sections [31]. Kwon et al.

Kwon et al. [32] utilized the high-speed camera to simultaneously collect melt pool images and location information under different laser powers. DNNs were applied to accurately classify melt pool images concerning different laser powers, as it resulted in different levels of porosity. Although SVMs and DNNs may detect anomalies through melt pool images in DED, the lack of label in real-time sensing limit the ability of these supervised learning methods. To tackle this problem, K-means clustering was employed to detect and locate such defects in an unsupervised manner. The K-mean disentangled data since intensity profiles of melt pool images pixels of overheated regions significantly differ from the normal melting state [33].

The K-mean disentangled data since intensity profiles of melt pool images pixels of overheated regions significantly differ from the normal melting state [33]. Self-organizing map (SOM) is also applied to thermal profiles of melt pools to identify abnormal melt pools and predict porosity [5].Although unsupervised learning is capable of differentiation of melt pool images based on the inherent structure in data, the lack of feature representation from ill-structured and noisy data limits380the capability to form a continuous and meaningful latent space. On the other hand, variational autoencoders have shown remarkable performance in unsupervised generative modeling of high- dimensional noisy data caused by complex distributions. VAEs as neural network-based models include two steps: encoding and decoding. The encoding step extracts nonlinear features automatically from original images to a low dimension matrix, which is called latent space.

The encoding step extracts nonlinear features automatically from original images to a low dimension matrix, which is called latent space. Then, extracted features could be converted to reconstructed images [34]. However, little has been done to leverage the sparse representation using VAEs for real-time defect detection in the DED process.Research MethodologyThis section presents the proposed methodology of sparse representation and unsupervised learning for defect monitoring using melt pool images in DED. As shown in Figure 1, our methodology is divided into the following steps: (1) melt pool image data acquisition, (2) synthetic data generation, (3) deep generative modeling, and (4) unsupervised learning of defects.Figure 1.

The flowchart of the proposed methodology for space representation and unsupervised learning of melt pool images in DED process.Here, the VAE consists of an encoder (i.e., a recognition network) that converts the input to a latent representation and a decoder (i.e., a generative network) that transforms the internal latent representation to the outputs.Part 1: Deep generative model for sparse representation:The essence of VAE is to learn a stochastic mapping between the input spaceempirical distribution and latent space considered as a generated model with a joint distributionwith simpler distribution. The latent space𝒙𝒙.with complex could be𝒛𝒛𝒛𝒛𝑝𝑝𝜽𝜽(𝒙𝒙, 𝒛𝒛)381(1) represent each thermal image and the is .where expression of 𝜽𝜽 a stochastic decoder, a new distribution of 𝑝𝑝𝜽𝜽(𝒛𝒛) 𝒙𝒙is a vector of the model’s parameter,and𝑝𝑝𝜃𝜃(𝒙𝒙, 𝒛𝒛) = 𝑝𝑝𝜽𝜽(𝑧𝑧)𝑝𝑝𝜽𝜽(𝒙𝒙|𝒛𝒛) 𝑧𝑧on latent space.

The less the KL-divergence is, the much similar the distributions of encoder and decoder are. The goal of VAE is to maximize . One of the most important properties of ELBO is that it allows us to use stochastic gradient descent (SGD) or Adam (i.e., an algorithm that combines properties of AdaGrad and RMSProp algorithms to handle sparse gradients on noisy problems) for joint optimization of encoder and decoder [35]. We random assign initial values of parameters ( ), and they will be stochastically optimized until convergence. To make the formulation tractable in VAE, we usually assume that follows Gaussian distribution:≤ 𝑙𝑙𝑙𝑙𝑙𝑙𝑝𝑝𝜽𝜽(𝒙𝒙)and to minimize the difference betweenand𝑙𝑙𝑙𝑙𝑙𝑙𝑝𝑝𝜽𝜽(𝒙𝒙)𝑞𝑞𝝓𝝓(𝒛𝒛|𝒙𝒙)𝑝𝑝𝜽𝜽(𝒛𝒛|𝒙𝒙)and𝝓𝝓𝜽𝜽𝑞𝑞𝝓𝝓(𝒛𝒛|𝑥𝑥) (8)2 are mean and standard deviation of latent space obtained from the encoder. Setting 𝑞𝑞𝝓𝝓(𝒛𝒛|𝒙𝒙) = 𝑁𝑁(𝝁𝝁, 𝝈𝝈 ) is the element-wise product.

The constructions of these two neural networks are set to be different. The architecture of encoder and decoder neural networks are represented in Figure 2. A rectified linear activation function (ReLu) is applied for activation functions of all hidden layers. Our model is optimized by Adam optimizer to obtain a better weight for the neural network using noisy manufacturing images.𝒙𝒙∈𝒟𝒟Figure 2. Neural networks architecture of variational autoencoder for thermal images.Part 2: Synthetic minority over-sampling techniqueThe unbalanced data in AM setting (i.e., low number of images representing defective regions compared to the whole number of images) prevents the VAE model from learning the decision boundary effectively. Synthetic Minority Over-sampling Technique (SMOTE), as one of the common methods, could generate minority cluster data to make the dataset balanced by oversampling the original dataset [36].

Synthetic Minority Over-sampling Technique (SMOTE), as one of the common methods, could generate minority cluster data to make the dataset balanced by oversampling the original dataset [36]. Therefore, we leverage SVM-SMOTE to generate some abnormal melt pool images. For SVM-SMOTE, the borderline area is identical to the support vectors, which are achieved by training traditional Support Vector Machine (SVM) on the original datasets [33]. New data could be created by order of the first nearest neighbor to the kth nearest neighbor instead of choosing the neighbors randomly. In general, minority cluster (i.e., a group of images portraying defective region in a build) is considered as positive and majority cluster (i.e., non-defective) is treated as negative. SVM-SMOTE trains the SVM on the inputs firstly to get the set of positive support vectors positive nearest neighbors nearest neighbors of each 𝑘𝑘and distribute input among them evenly. Combineto get an array as 𝑺𝑺𝑺𝑺 +.

Combineto get an array as 𝑺𝑺𝑺𝑺 +. After that, if half of+𝑺𝑺𝑺𝑺𝒏𝒏𝒏𝒏𝑚𝑚383(10)(11)(12)is the number of positive support vectors, , where come from negative cluster, then is considered as an array contains the number of generated data corresponding to each positive 𝐴𝐴𝐴𝐴 = {𝑎𝑎𝑎𝑎𝑡𝑡}𝑡𝑡=1 with its k positive new positive class data among the lines joining nearest neighbors. The generated melt pool images can be represented as 𝑆𝑆𝑆𝑆𝑇𝑇𝑇𝑇.

The generated melt pool images can be represented as 𝑆𝑆𝑆𝑆𝑇𝑇𝑇𝑇. Next, we create+ 𝒔𝒔𝒔𝒔𝑖𝑖𝑎𝑎𝑎𝑎𝒊𝒊(13)+ = 𝒔𝒔𝒔𝒔𝑖𝑖+ + 𝜌𝜌 × (𝒔𝒔𝒔𝒔𝑖𝑖+ 𝒙𝒙𝑛𝑛𝑛𝑛𝑛𝑛is the jth minority nearest neighbor ofwhere and generated data can be expressed by− 𝒏𝒏𝒏𝒏𝑖𝑖,𝑗𝑗)𝒏𝒏𝒏𝒏𝑖𝑖,𝑗𝑗𝜌𝜌 ∈ [0,1]+ 𝒔𝒔𝒔𝒔𝑖𝑖.On the contrary, the(14)+ = 𝒔𝒔𝒔𝒔𝑖𝑖+ + 𝜌𝜌 × (𝒏𝒏𝒏𝒏𝑖𝑖,𝑗𝑗 − 𝒔𝒔𝒔𝒔𝑖𝑖 )+ 𝒙𝒙𝑛𝑛𝑛𝑛𝑛𝑛Finally, we could get balanced data by combining the original data and generated data.Part 3: T-distributed stochastic neighbor embedding dimension reductionT-distributed stochastic neighbor embedding (t-SNE) is one of the statistical models which maps high-dimensional data to two or three dimensions for each data point [37]. It is a nonlinear dimensionality reduction algorithm, which is suitable for visualizing high-dimensional data by embedding it into lower dimensions. t-SNE converts the distance between the data points in the original space into a Gaussian distribution probability.

t-SNE converts the distance between the data points in the original space into a Gaussian distribution probability. The closer the distance in high-dimensional space is, the larger the probability. In our case, for vitalizing location of images after mapping them into latent space more clearly. t-SNE is used for reducing the dimension of latent variables from N to 2. It computes the proportional to the similarity of objects:𝑥𝑥𝑖𝑖, 𝑥𝑥𝑗𝑗2 exp (−𝑑𝑑(𝒛𝒛𝑖𝑖, 𝒛𝒛𝑗𝑗)/2𝜎𝜎𝑖𝑖 ) is the latent variables of a melt pool image, 2 𝑘𝑘≠𝑖𝑖 ∑ exp (−𝑑𝑑(𝒛𝒛𝑖𝑖, 𝒛𝒛𝑗𝑗)/2𝜎𝜎𝑖𝑖 , which corresponds to :where 2D, we could get 𝒛𝒛𝑖𝑖 represent the similarity of𝑝𝑝𝑗𝑗|𝑖𝑖 =, 𝑝𝑝𝑖𝑖|𝑖𝑖 = 0 and. After mapping) 𝑖𝑖, 𝑗𝑗 ∈ 𝐿𝐿,. Eq.

After mapping) 𝑖𝑖, 𝑗𝑗 ∈ 𝐿𝐿,. Eq. (16) leverages Student-t Distribution to 𝒛𝒛𝑖𝑖 𝑖𝑖 ≠ 𝑗𝑗(15)into𝑦𝑦𝑖𝑖𝑧𝑧𝑖𝑖𝑦𝑦𝑖𝑖, 𝑦𝑦𝑗𝑗2−1(1 + ||𝑦𝑦𝑖𝑖 − 𝑦𝑦𝑗𝑗|| ) 𝑘𝑘≠𝑙𝑙 ∑ (1 + ||𝑦𝑦𝑘𝑘 − 𝑦𝑦𝑙𝑙|| 𝑦𝑦𝑖𝑖where function as following:Then the location ofis determined by optimizing the KL divergence and the loss𝑞𝑞𝑖𝑖𝑗𝑗 =, 𝑞𝑞𝑖𝑖|𝑖𝑖 = 02−1)𝑖𝑖 ≠ 𝑗𝑗.(16)𝑝𝑝𝑖𝑖𝑗𝑗 𝑞𝑞𝑖𝑖𝑗𝑗𝐾𝐾𝐿𝐿(𝑃𝑃|𝑄𝑄) = � 𝑝𝑝𝑖𝑖𝑗𝑗𝑙𝑙𝑙𝑙𝑙𝑙A gradient descent algorithm is used for minimizing KL divergence. Here,𝑖𝑖≠𝑗𝑗to the result of minimization.𝑦𝑦𝑖𝑖(17)is correspondingPart 4: Clustering modelsIn this part, we are using two different methods to do the clustering: Gaussian Mixture Model (GMM) and K-Means. We are interested in finding out whether the latent variables of VAE explain their cluster. Therefore, the constructed latent values are the inputs of these two clustering methods. Both for K-Means and GMM are trained with an assumption that the number of clustering is 2.

Therefore, the constructed latent values are the inputs of these two clustering methods. Both for K-Means and GMM are trained with an assumption that the number of clustering is 2. K-Means:K-Means is an iterative clustering algorithm based on squared Euclidean distance. This method aims to assign N observations into k clusters in which every observation belongs to a cluster with384the closest squared Euclidean distance. Here, the clustering, and the algorithm proceeds as follows:. In our study, we use standard K-Means to do1. Choose a set of k means as initial cluster centroids 2. For each latent space point, calculate Euclidean distance to every centroid and assign each 𝑘𝑘 = 2(𝑡𝑡)(𝑡𝑡)= {𝒛𝒛𝑝𝑝: �𝒛𝒛𝑝𝑝 − 𝒎𝒎𝑖𝑖� ≤ �𝒛𝒛𝑝𝑝 − 𝒎𝒎𝑗𝑗𝑺𝑺𝑖𝑖3. Recalculate the centroids of each cluster.(𝑡𝑡)�2∀𝑗𝑗, 1 ≤ 𝑗𝑗 ≤ 𝑘𝑘}(18)(𝑡𝑡+1)(19)𝒎𝒎𝑗𝑗=4. Repeat steps 2 and 3 until the convergence. (𝑡𝑡)4. Repeat steps 2 and 3 until the convergence.

Repeat steps 2 and 3 until the convergence. (𝑡𝑡)4. Repeat steps 2 and 3 until the convergence. (𝑡𝑡)Gaussian mixture model:GMM is an extension of a single Gaussian probability density function. The model is comprised of several different single Gaussians. Each Gaussian is regarded as a cluster according to the different parameters of the normal probability density function (PDF). By computing the PDF of the input, the typical cluster of the input could be determined by a threshold. The PDF of inputs could be represented as Eq. (20):𝑀𝑀(20)where is the weight factor and denotes the probability of choosing the kth Gaussian model. With the sample points with ) could be calculated.

An OPTOMEC Lens 750 machine equipped with a 1-kW Nd: YAG laser, pyrometer, and an in- chamber thermal camera (Figure 3(a)) are utilized to fabricate single-track Ti-6AI-4V thin walls (Figure 3(b)). The experiment was performed at Mississippi State University with popular Lens technique in DED process that has the capability to create functionally graded materials [5]. To predict the porosity occurrence and the time-varying evolution of melt pool dynamics, a built-in thermal imaging system is integrated. Temperature distribution of the top surface of melt pool during fabrication is recorded by a dual-wavelength pyrometer (Stratonics, Inc.). Due to the specified exposure time (2.0274 ) for a given collection rate of the pyrometer, the sensor could reduce the chance of blur resulting from motion compared to the IR camera. In this experiment, in each exposure time.

In this experiment, in each exposure time. There the laser’s scan rate and build move are set to 12.7 for each pixel is a CMOS detector applied in the pyrometer with𝑚𝑚𝑚𝑚and 26pixels and 6.45𝑚𝑚𝑚𝑚/𝑚𝑚𝜇𝜇𝑚𝑚752 × 480𝜇𝜇𝑚𝑚385size. The temperature in the pyrometer has a range of 1000 to 2500 clock to 2.0274and 5, respectively.. The exposure time and pixel℃𝑚𝑚𝑚𝑚𝑀𝑀𝑀𝑀𝑧𝑧Figure 3. (a) The advanced imaging system installed in the OPTOMEC LENS 750 machine, and (b) final thin wall builds using Ti-6Al-4V material [5].Because of the high image resolution (1.7for each melt pool image) and monitoring , thickness =frequency, the build of a single thin wall (length = 47.81 1.78, height = 27.56𝑀𝑀𝑀𝑀) brings about 4.7of image stream data.𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝑚𝐺𝐺𝑀𝑀Figure 4.

(a) The image of the thin wall, and (b) The relationship between meltpool characteristics and lack of fusion.Figure 4 demonstrates the potential relationship between characteristics of melt pool images and lack of fusion from our experimental data. Note that the color of meltpool indicates the temprature of the melp pool. After obtaining the data, a radon transform is applied to reconstruct the image using the acquired X-ray intensity readings for each of the detector elements. Then the individual slice images could be accumulated into a 3D reconstruction. Next, a 3D reconstruction could be generated by accumulation of the individual slice images. The porosity of the as-built part, which is derived from melt pool characteristics, is compared with the XCT characterization. The thermal image data are processed and stored in1564 thermal image files, each of which contains 480×752 cells of information. The area of the melt pool is extracted based on the melt =1636◦C).

The thermal image data are processed and stored in1564 thermal image files, each of which contains 480×752 cells of information. The area of the melt pool is extracted based on the melt =1636◦C). Each melt pool is scaled using the coordinates of the pool temperature of Ti-6Al-4V( peak temperature. The accuracy of prediction is examined by finding the proportion of pores, which are predicted correctly using the melt pool clustering method, in total pores.𝑇𝑇𝛾𝛾386Experimental ResultsIn this section, we perform the experimental study on our dataset with the proposed methodology. To prevent overfitting, we perform K-fold cross validation with different fold numbers. Note that we have 1493 positive instances (i.e., defective images) and 71 (i.e., non- defective images) negative instances in our data. We first performed SMOTE-SVM to balance the positive and negative cases for VAE construction; therefore, we have in total 2896 instances.

We first performed SMOTE-SVM to balance the positive and negative cases for VAE construction; therefore, we have in total 2896 instances. We trained the VAE based on the proposed methodology and run each experiment 20 times. Note that each image has a size of 28 × 28. We investigate VAE-Kmeans and VAE-GMM with different dimensions of latent space (L=2, 5, 10, 100, 1000) and the number of folds (K = 2, 3, 4, 5).We first utilize the t-SNE to visualize the embedded latent space of the whole dataset. As shown in Figure 4, the trend of separation can be seen from a) to e). Each node in the t-SNE visualization is corresponding to one image in the dataset.Figure 4.

As shown in Figure 4, the trend of separation can be seen from a) to e). Each node in the t-SNE visualization is corresponding to one image in the dataset.Figure 4. t-SNE embedding of the latent space based on the number of latent in VAE, a) number of latent L=2, b) number of latent L=5, and c) number of latent L=10, d) number of latent L=100, and e) number of latent L=1000.Figure 4 a) shows the nodes from two classes are overlapped with each other, indicating that the two-dimensional latent space embedding cannot extract enough useful information to express the distinction of two classes. However, with the increment of the number of latent L, the distinction between two groups can be clearly visualized through t-SNE embedding. In Figure 4 (e), only a few images are misclassified into the other group.Figure 5.

In Figure 4 (e), only a few images are misclassified into the other group.Figure 5. The accuracy of porosity detection using the proposed VAE-KMeans and VAE-GMM when embedding to different number of latent for a) 2 folds, b) 3 folds, c) 4 folds, and d) 5 folds.Figure 5 illustrates the accuracy of porosity detection using the proposed VAE-KMeans and VAE-GMM when embedding to different numbers of latent for a) 2 folds, b) 3 folds, c) 4 folds, and d) 5 folds. As shown in Figure 5, the overall trend of detection accuracy increases when the number of latent increases. However, the number of latent that brings the best detection accuracy under different fold numbers varies. For example, the best choice for the number of latent L=10 while K=2, and L=100 is a better option when K=2 since under different values of K, the number387of training data and testing data are different. The training set has more data when K=2 in comparison with K=3.

The training set has more data when K=2 in comparison with K=3. It can be seen that less amount of training data needs more dimensions in latent to have a better detection accuracy. Also, it is worth mentioning that the accuracies under L=1000 among the four subplots in Figure 5 are not the highest, indicating that a higher dimension in latent space does not necessarily mean that it can capture all the hidden information from melt pool images and provide a better detection result. Instead, it might pick redundant features from the images. As shown in Figure 5 (c) and Figure 5 (d), the accuracies under different L could reach 0.9. Even though the accuracies whenare the highest among all experiments when, which means it performs better 𝐾𝐾 = and the computational cost is lower. Overall, K=4 results in a better porosity detection than 5 accuracy with the consideration of the computational complexity.

Overall, K=4 results in a better porosity detection than 5 accuracy with the consideration of the computational complexity. Further, we construct the confusion matrix and report the sensitivity (SEN), specificity (SPC), precision (PPV), and F-scores, respectively, under, all accuracies under different L are higher than 0.7 when𝐿𝐿 = 100𝐾𝐾 = 4𝐾𝐾 = 5. The results of the experiment are provided in Table 1.𝐾𝐾 = 4Table 1.

Interestingly, when , the values of F-scores are highest both for K-Means and GMM, resulting in 0.885±0.0156 and 0.871±0.0172, which indicates the highest accuracies in terms of porosity prediction.𝐹𝐹 − 𝑚𝑚𝑆𝑆𝑙𝑙𝑃𝑃𝑆𝑆 =𝐿𝐿 = 100(23)(24)(25)ConclusionIn-situ thermal image sensing systems are recently developed to address the quality challenge in DED. However, they are still at an early age and are limited in the ability to analyze generated data in real-time. The presence of high-dimensional, low signal-to-noise ratio, and ill-structured data without annotation hinders real-time defect detection in the process. There is an urgent need to integrate real-time thermal image data with newly developed machine learning methods and realize the qualify-as-you-build paradigm in DED. We validate the proposed porosity prediction388method by benchmarking against the X-ray CT characterization of a Ti-6Al-4V thin wall build fabricated using a LENS system.

We validate the proposed porosity prediction388method by benchmarking against the X-ray CT characterization of a Ti-6Al-4V thin wall build fabricated using a LENS system. The comparative study with a previous work that focuses on the morphological features of melt pools (e.g., area, length, width, etc.) shows significant improvement in prediction accuracy. The clustering based on morphological features fails to form distinct clusters, which indicates that simple features are not sufficient for the identification of anomalies. The structure of melt pool images can be learned through the sparse representation model that encodes both morphological features and thermal distribution automatically.AcknowledgmentThe authors gratefully acknowledge the valuable contributions from the Center for Advanced Vehicular Systems (CAVS) at Mississippi State University for this research.Reference[1] A. Gaikwad, R. Yavari, M. Montazeri, K. Cole, L. Bian and P.

M. Colosimo and M. Grasso, "On-Machine Measurement, Monitoring and Control,"Precision Metal Additive Manufacturing, pp. 102, 2020.[20] Q. Tian, S. Guo, E. Melder, L. Bian and W. Guo, "Deep Learning-Based Data FusionMethod for In Situ Porosity Detection in Laser-Based Additive Manufacturing," Journal of Manufacturing Science and Engineering, vol. 143, (4), p. 041011, 2021.[21] S. A. Shevchik, C. Kenel, C. Leinenbach and K. Wasmer, "Acoustic emission for in situquality monitoring in additive manufacturing using spectral convolutional neural networks," Additive Manufacturing, vol. 21, pp. 598-604, 2018.[22] H. Wu, Z. Yu and Y. Wang, "A new approach for online monitoring of additive manufacturing based on acoustic emission," in ASME 2016 11th International Manufacturing Science and Engineering Conference, 2016, p. V003T08A013.[23] J. C. Heigel, P. Michaleris and E. W.

V003T08A013.[23] J. C. Heigel, P. Michaleris and E. W. Reutzel, "Thermo-mechanical model development and validation of directed energy deposition additive manufacturing of Ti–6Al–4V," Additive Manufacturing, vol. 5, pp. 9-19, 2015.[24] L. Song, V. Bagavath-Singh, B. Dutta and J. Mazumder, "Control of melt pool temperature and deposition height during direct metal deposition process," The International Journal of Advanced Manufacturing Technology, vol. 58, (1), pp. 247-256, 2012.[25] D. A. Kriczky, J. Irwin, E. W. Reutzel, P. Michaleris, A. R. Nassar and J. Craig, "3D spatial reconstruction of thermal characteristics in directed energy deposition through optical thermal imaging," J. Mater. Process. Technol., vol. 221, pp. 172-186, 2015.390[26] J. P. Davim, C. Oliveira and A. Cardoso, "Predicting the geometric form of clad in laser cladding by powder using multiple regression analysis (MRA)," Mater Des, vol. 29, (2), pp. 554-557, 2008.[27] A. R. Nassar, T. J. Spurgeon and E.

TOMATED CONSTRUCTION USING CONTOUR CRAFTINGB. Khoshnevis, H. Kwon, and S. BukkapatnamIndustrial and Systems Engineering, University of Southern California Los Angeles, CA 90089 khoshnev@usc.eduAbstractThis paper presents some concepts and initial investigation of a novel construction automation approach using the Contour Crafting (CC) layered fabrication process, developed at the University of Southern California. CC uses computer control to take advantage of the superior surface forming capability of trowels, used by craftsmen and builders since ancient times, to create large intricate structures with smooth and accurate surfaces. The potential of CC became evident from the initial investigations and experiments with various materials and geometries.

The potential of CC became evident from the initial investigations and experiments with various materials and geometries. Using this process, a single house or a colony of houses, each with possibly a different design, may be automatically constructed in a single setup.IntroductionPresently, the construction industry is facing various problems including high project costs, low labor efficiency, high at-site accident rates, vanishing skilled workforce, and poor control of construction projects (Warszawski, and Navon, 1998). Automation has resolved several similar problems in the manufacturing industry. The construction industry, however, largely remains manual and manpower intensive. The few attempts made towards automating certain aspects of construction have aimed only at mechanizing the same manual approach (e.g., use of a brick-laying robot) without introducing any new paradigm (Pegna, 1997).

Development of new automation paradigms for whole structure construction may mitigate many of the problems that the industry is facing. In fact, new automation paradigms seem imperative for several applications, including construction of emergency homes and low income housing projects. Furthermore, the development of construction automation technologies is necessary if colonization of other planets is to become a reality in the coming century.Conventional methods of manufacturing automation do not lend themselves to construction of large structures with internal features. This explains why the evolution of construction automation has been slow. A promising new automation approach is layered fabrication, generally known as solid free form fabrication or rapid prototyping, which uses an additive method and is capable of creating complex internal features.

However, most of the current layered fabrication methods are limited by their ability to deliver a wide variety of materials applicable to construction. Additionally, they are severely constrained by the low rates of material deposition that makes them attractive only to the fabrication of small industrial parts.497Contour Crafting (CC) is a layered fabrication technology that uses computer control to exploit the superior surface-forming capability of troweling to create smooth and accurate planar and free-form surfaces (Khoshnevis et al., 2001; Khoshnevis, 1998). Using the layering approach afforded by troweling, an ancient process, a wide range of surface shapes may be created with fewer types of troweling tools than are needed for traditional plaster handwork and sculpting.

Some of the important advantages of CC compared with other layered fabrication processes are better surface quality, higher fabrication speed, and a wider choice of materials.The key feature of CC is the use of two trowels, which in effect act as two solid planar surfaces, to create surfaces on the object being fabricated that are exceptionally smooth and accurate. Artists and craftsmen have effectively used simple tools such as trowels, blades, sculpturing knives, and putty knives, shown in Figure 1, with one or two planar surfaces for forming materials in paste form since ancient times. Their versatility and effectiveness for fabricating complex free-form as well as planar surfaces is evidenced by ancient ceramic containers and sculptures with intricate or complex surface geometries as well as detailed plaster work that have shapes as complicated as flowers, on the walls of rooms.Figure 1.

Simple Historical Construction ToolsIn CC, computer control is used to take advantage of the superior surface forming capability of troweling to create smooth and accurate, planar and free-form surfaces (Landfoam FAQ, 2000). The layering approach enables the creation of various surface shapes using fewer different troweling tools than in traditional plaster handwork and sculpting. It is a hybrid method that combines an extrusion process for forming the object surfaces and a filling process (pouring or injection) to build the object core. As shown in Figure 2, the extrusion nozzle has a top and a side trowel. As the material is extruded, the traversal of the trowels creates smooth outer and top surfaces on the layer. The side trowel can be deflected to create non-orthogonal surfaces. The extrusion process builds only the outside edges (rims) of each layer of the object.

The side trowel can be deflected to create non-orthogonal surfaces. The extrusion process builds only the outside edges (rims) of each layer of the object. After complete extrusion of each closed section of a given layer, if needed filler material such as concrete can be poured to fill the area defined by the extruded rims.Material feed barrelSide trowel control mechanismNozzleTop trowelSide trowelFigure 2. Contour Crafting ProcessSome internal walls can be extruded within each layer to create square hatches or other types of hatches (see Figure 3). The hatching process may be required for large objects, since setting or curing can start before the filler material gets a chance to spread over the entire surface of the layer.498However, when hatching is used, each of the small hatches is filled separately, which because of their small size allows more control over the spreading and curing of the filler material.

Hatching can also accelerate the forming process because it provides for concurrent extrusion and filling (i.e., as the extrusion nozzle creates new hatches, previously made hatches can be filled).Advantages of CC for automated constructionFigure 3. Creation of Internal WallsAmong all things made by man building structures are those that are fabricated in a layered fashion. This is a clear indication that layered fabrication could be the ideal approach for automating building construction.

This is a clear indication that layered fabrication could be the ideal approach for automating building construction. Although several methods of layered fabrication have been developed in the last two decades (Pegna, 1997), and successful applications of these methods have been reported in a large variety of domains (including industrial tooling, medical, toy making, etc.), only Contour Crafting (CC) is uniquely applicable to construction of large such as houses structures (Khoshnevis, 2000).The immediate impact of CC will be in the construction of emergency and low income housing structures in various regions by taking advantage of the variety of locally available resources including soil, gravel, lime stone, clay, wood chips, etc.

The applicability of CC can possibly be extended to construction applications in regular residential and commercial structures that could utilize functional and exotic architectural geometries, which are difficult to realize using the current m a n u a l p r a c t i c e . c o n s t r u c t i o n Furthermore, building of structures on other planets using local materials may be achieved using the proposed technology. Such application in building construction is depicted in Figures 4 where a gantry system carrying the CC nozzle moves on two parallel lanes installed at the construction site. A single house or a colony of houses, each with possibly a different design, may be automatically constructed in a single run.

A single house or a colony of houses, each with possibly a different design, may be automatically constructed in a single run. Following are some interesting aspects of this automated construction process:Figure 4.Residential Building Construction499(cid:137) The process allows architects to design structures with functional and exotic architectural geometries that are difficult to realize using the current manual construction practice. The surfaces produced will be paint-ready, hence an automated painting system may be integrated into the CC system.(cid:137) Since deposition in CC is controlled by computer, accurate amounts of selected construction materials, such as smart concrete, may be deposited precisely in the intended locations. This way the electric resistance, for example, of a carbon filled concrete may be accurately set as dictated by the design.

This way the electric resistance, for example, of a carbon filled concrete may be accurately set as dictated by the design. Elements such as strain sensors, floor and wall heaters can be built into the structure in an integrated and fully automated manner.Figure 5. Complex Wall Section(cid:137) As shown in Figure 5 utility conduits may be built into the walls of a building structureprecisely as dictated by the CAD data.(cid:137) The quality of surface finish in CC is controlled by the trowel surface and is independent ofthe size of the nozzle orifice. Consequently, various additives such as sand, gravel, reinforcement fiber, and other applicable materials available locally may be mixed and extruded through the CC nozzle.(cid:137) Multiple materials that chemically react may be fed through the CC nozzle system and mixed in the nozzle barrel immediately before deposition.

The quantity of each material may be controlled by computer and correlated to various regions of the geometry of the structure being built. This will make possible the construction of structures that contain varying amounts of different compounds in different regions.(cid:137) Methods of imbedding of steel mesh and other forms of reinforcement into each layer may bedevised. Note that in this configuration the CC nozzle, the inter-rim filler nozzle, and reinforcement placement mechanisms can all be on the same gantry system. The proposed system can create shapes with smooth outer surfaces and reinforced internal structure automatically and in one setup.(cid:137) The lower construction cost and time that CC offers will make frequent reconstruction of buildings a more feasible proposition. This is of special interest in commercial building sector.

This is of special interest in commercial building sector. This possibility is especially appealing if recyclable materials are used.Experimental investigationsWe conducted extensive experiments which show the potential of CC for construction automation. Many of the advantages mentioned in the previous section have been validated through these investigations.Material used for the experimentsOur examination of the properties of a few ceramic materials and compounds revealed that a type of clay procured from America Ware Co. in Los Angeles did not swell upon exiting500the nozzle orifice. This probably was due to its readiness to shear and virtual absence of an elastic phase (Bardet, 1997; Craig, 1997). We qualified the Bingham characteristics of this material (Fox and McDonald, 1985; Levy and Carley, 1994).

We qualified the Bingham characteristics of this material (Fox and McDonald, 1985; Levy and Carley, 1994). Using this material we have been able to create the several geometrical features on our CC machine.Description of CC machineThe machine shown in Figure 6 is modified to fabricate the complex geometrical parts for the CC process of uncured ceramic materials, which include the convex, concave, and sharp corner shapes. The machine consists of a trowel rotation system, and a vertical extrusion head capable of linear motion along three coordinate axes.Figure 6. The CC machine for fabricating complex geometriesThe trowel rotation mechanism shown in Figure7 consists of a bevel gear, and a connector. The ratio of the bevel gear is 4 to 1, and is derived by the 5th stepper motor.

The ratio of the bevel gear is 4 to 1, and is derived by the 5th stepper motor. The connection mechanism allows the raw material to flow continuously from the cylinder to nozzle, and can rotate the extrusion system without disturbing the material flow while fabricating complex curves.Cylinder ConnectorBevel gearNozzleSide trowelThe extrusion system consists of a top and side trowel, a cylinder that contains the raw material, and a piston and a threaded feed rod that extrudes the raw material through a nozzle. The process utilizes a Programmable Multi-Axis Controller (PMAC), a high- performance servo motion controller, capable of controlling up to eight axes of motion (Delta Tau Data Systems, 1996 a, b). The eight axes can be all synchronized for completely coordinated motion; each axis can be put into its own coordinate system for eight completely independent operations; any intermediate arrangement of axes into coordinate systems is also possible.

Limit switches are used to restrict motion to specified limits.Figure 7. Trowel rotation system of the CC processControl mechan of movable side trowelIn order to enhance the capability of CC for building certain primitive geometries and hybrid geometries made of these primitives, a nozzle with movable side trowel is designed and assembled into the existing machine as depicted in Figure 8. This nozzle assembly provides for intricate trowel motions to create various complex geometries.Figure 8.Side view of the movable nozzle assembly501Experimental resultsFabricating complex 3D shapesAdapting the movable side trowel control mechanism shown in Figure 9 (a), our existing CC fabrication machine has been modified for fabricating a variety of 2.5D and 3D shapes that constitute the primitive construction geometries (e.g. flat floors, straight walls, pyramid, domes, etc.) and hybrid geometries made of these primitives.

flat floors, straight walls, pyramid, domes, etc.) and hybrid geometries made of these primitives. The primitives have been carefully chosen so that these form the basic shapes for the scaled models of adobe houses (Khalili, 2000). As shown in Figure 9 (b), typical primitives that we have considered thus far are flat floors, straight walls, pyramid, and domes. Hybrid geometries, formed as a direct combination of these primitives, will represent scaled models of adobe houses such as those adopted by CalEarth in Figure 14.Figure 9. CC process with the movable side trowel: and various primitive partsRoofs, overhangs, and support structuresThrough extensive experimental investigations for the support material, Wax and sand were used as the support materials in order to construct overhang features such as roofs as shown in Figure 10. Note that although this approach is feasible for small parts it may not be so for housing structures.

Note that although this approach is feasible for small parts it may not be so for housing structures. We consider using other means of building overhangs that are more suitable for constructing such structures (one such approach is depicted in Figure 16 for roofing adobe houses).Depositions with hollow cavitiesIn order to investigate co-extrusion of multipleFigure 10. Fabrication of roofs, overhangs, and support structuresmaterials using a combination of nozzles, extensive experimentation and a series of design enhancements have been conducted. As shown in Figure 11, we have developed the capability to build layers with hollow depositions using CC. Mandrel of various shapes may be used for creating hollows of various shapes.With these cavities by forcing the material through a nozzle with a central mandrel, various materials may be co-extruded if the mandrel itself is hollow and works as a nozzle to deliver a second material. ThisFigure 11.

ThisFigure 11. Laying hollow sections through CC: (a) the ceramic material in the nozzle before extrusion, (b) hollow circle formed as the extrudate emerges from the orifice, and (c) cross section of the fabricated part revealing the hollow sections502feature will provide the capability to lay base material as well as certain reinforcements simultaneously. These hollow cavities also result in lighter structures.Furthermore, for constructing walled structures, we will first construct rims using the nozzle system, and will fill the intervening space between wall-rims using a bulk-filling mechanism, a schematic of which is shown in Figure 3.Reinforcements and impregnationTowards improving the strength of large housing structures through CC, we have built investigated the use of a variety of reinforcements. For example, Figure 12 shows pictures from our c o i l e x p e r i m e n t s reinforcement.

For example, Figure 12 shows pictures from our c o i l e x p e r i m e n t s reinforcement. Owing to the high extrusion pressures prevailing in CC compared to other layered free-form fabrication techniques (Zak et al., 1999), the extrudate thoroughly adheres itself around the coils without causing any internal discontinuities. Similar results have been observed from our experiments with sand impregnation.Figure 12. Reinforcement process of CC: (a) metal coil placed on a top layer, (b) a fresh layer of extrudate covers the coil, and (c) cross sections of the fabricated part with the reinforcement coil showing a reasonable adhesion between layersw i t hAs shown in Figure 13, two simple modular components may be delivered by an automated feeding system that deposits and assembles them between the two rims of each layer built by CC. Concrete may then be poured between the rims of each layer to contain the steel mesh. The mesh can follow the geometry of the structure.

Concrete may then be poured between the rims of each layer to contain the steel mesh. The mesh can follow the geometry of the structure. Note that in this configuration the CC nozzle, the steel feeder, and the concrete filler feeder are all on the same gantry system. Thus the use of reinforcements seems promising in our CC process.Figure 13. Steel reinforcement modules and meshSummary and future workPreliminary investigation of CC process indicates that the process is feasible and has significant potential in construction automation. The potential of CC became evident from the presented experimental investigations. Experiments with ceramics material show the versatility of the process relative to the use of a variety of fabrication materials.

Experiments with ceramics material show the versatility of the process relative to the use of a variety of fabrication materials. The process seems especially suited for automated construction of emergency structures as well as large scale construction of houses with exotic features such as adobe houses.In order to facilitate accurate deposition of material and efficient fabrication of the proposed housing structures, we are currently developing a computer control software by modifying our current control software for fabricating both primitive as well as hybrid geometries. The new software will allow for experimenting with various types of trowel motions503and material flow patterns. In devising our construction control software we will benefit from the ancient body of knowledge that is currently being harnessed by CalEarth for building supportless closed structures.

In devising our construction control software we will benefit from the ancient body of knowledge that is currently being harnessed by CalEarth for building supportless closed structures. An example of a clever and ancient manual method of constructing such supportless structures is shown in Figure 14. Here the vault is constructed by first laying side squinches, and the dome structure is built layer-by-layer. Figure 15 shows a vault structure made of clay bricks using this traditional manual procedure. Our corresponding deposition pattern, inspired by these ancient methods, could be such as the one schematically illustrated in Figure 16.Figure 14. Manual construction of adobe form structures using clay bricks (Source: Khalili, 2000)Figure 15. A vault structure made of clay bricks (Source: Khalili, 2000)Figure 16.

TOMATED DESIGN OF TISSUE ENGINEERING SCAFFOLDS BY ADVANCED CAD E. Ramin* and R. A. Harris* * Rapid Manufacturing Research Group, Wolfson School of Mechanical & Manufacturing Engineering, Loughborough University, UKAbstractThe design of scaffolds with an intricate and controlled internal structure represents a challenge for Tissue Engineering. Several scaffold manufacturing techniques allow the creation of complex and random architectures, but have little or no control over geometrical parameters such as pore size, shape and interconnectivity- things that are essential for tissue regeneration. The combined use of CAD software and layer manufacturing techniques allow a high degree of control over those parameters, resulting in reproducible geometrical architectures.

The combined use of CAD software and layer manufacturing techniques allow a high degree of control over those parameters, resulting in reproducible geometrical architectures. However, the design of the complex and intricate network of channels that are required in conventional CAD, is extremely time consuming: manually setting thousands of different geometrical parameters may require several days in which to design the individual scaffold structures. This research proposes an automated design methodology in order to overcome those limitations. The combined use of Object Oriented Programming and advanced CAD software, allows the rapid generation of thousands of different geometrical elements. Each has a different set of parameters that can be changed by the software, either randomly or according to a given mathematical formula, so that they match the different distribution of geometrical elements such as pore size and pore interconnectivity.

This work describes a methodology that has been used to design five cubic scaffolds with pore size ranging from about 200 to 800 µm, each with an increased complexity of the internal geometry.Keywords: Tissue Engineering scaffolds, Automated Design, Advanced CAD Programming, Layer Manufacturing1. IntroductionThe goal of Tissue Engineering (TE) is to "restore function through the delivery of living elements which become integrated into the patient” (Vacanti and Langer, 1999). Currently, there are two areas of investigation for the physical realisation of TE. The first, termed 'direct organ printing', approaches the direct manufacture of organs and tissues using Layer Manufacturing (LM) methodologies, adapted in order to process biological materials, cells and living tissues (Mironov et al., 2003). It will probably be several years before the direct printing of living tissues becomes a reality.

It will probably be several years before the direct printing of living tissues becomes a reality. The second approach consists of producing scaffolds which are latterly seeded with cells. These scaffolds feature macro and micro porosity to promote directed natural tissue regeneration in vitro. An alternative approach involves the implantation of scaffolds without cells directly into the defect cavity in vivo, to serve as a template for cell and tissue growth (Capito and Spector, 2003). In vitro tests show that the lack of oxygen and nutrients affects the depth of penetration inside the scaffold, which, for foam structures, is less than 500 micron. That has been attributed to limited control over pore size and interconnectivity. Therefore, cells that colonise theperiphery of the scaffold, act as a barrier for a deeper diffusion of oxygen and nutrients, thus preventing new cells migrating from further (Sachlos and Czernuszka, 2003).

Traditional TE scaffold manufacturing processes attempt to control the micro-architecture of a scaffold, using a variation of the process parameters. These techniques are limited in their resolution and reproducibility. They might generate a wide distribution of pore sizes but they have poor control over interconnection, geometry, and spatial distribution- things that are essential for the production of regenerated tissue (Taboas et al., 2003). The use of LM techniques to create TE scaffolds has many advantages. Such technologies allow the realisation of implants with customized external shape via combination with reverse engineered CT or MRI data. Reproducible internal structures are obtainable, giving full control over distribution, interconnection, porosity and channel geometry. That enables an increased and controlled flow inside the scaffold, and therefore aids the regeneration of new tissue.

That enables an increased and controlled flow inside the scaffold, and therefore aids the regeneration of new tissue. In addition they are cost effective since their manufacture does not involve the production of tooling—thereby allowing the realisation of customised implants.1.1 Bone StructureThe optimum geometry for TE scaffolds for bone should be similar to that of the native tissue in need of repair. (Lin et al., 2004). TE scaffolds should mimic the structures of both the trabecular (spongy) and cortical (dense) sections of bone. In order to assess and analyse these microstructures, several studies have been conducted using non-destructive techniques--such as Micro-Computed Tomography (µCT)--that are capable of acquiring images of small bone samples with resolutions of 1 µm (Genant et al., 1999). Jones et al., (2004) used µCT to analyze the morphology and interconnectivity of the regenerated tissue into a polymer scaffold.

In particular they need: (1) high porosity and interconnecting pores of appropriate size and curvature. The optimal architecture should mimic that of the original tissue; (2) bio-materials with controlled degradation and resorbability rates; (3) adequate mechanical properties for load-bearing applications; (4) a tailored external shape to fit anatomical defects and (5) they must be easily to manufacture and have reproducible internal structures. While there have been many works on the physical aspects of these requirements, there has been comparatively little investment that considers the design requirements of these highly complex geometries.Most previous research for the design of TE scaffolds for LM has relied on libraries of unit cells, each with a different, predefined and generally regular internal architecture that does not mimic the natural, highly irregular, structure of bones.

Furthermore, the use of unit cells allows control only over the size of some geometrical elements, not their shape, which leads to more regular structures. The selected unit cell is then repeated in 3D to form a prism that contains the entire implant. Using Boolean operations, these two geometries are then combined to design a scaffold with the external shape of the implantation site and the predefined internal structure. Interesting examples of similar approaches can be found in Chua et al. (2003), Das and Hollister (2003), Naing et al. (2005), Hutmacher et al. (2001), Rosen et al. (2006).1.3 Advantages of Automated CAD DesignThe aim of this work was to focus attention on an efficient CAD methodology, in order to rapidly enable the design and integration of an intricate network of channels as dictated by a set of variable parameters. The input data can be changed either randomly, or according to given mathematical functions, leading to infinite different networks.

The input data can be changed either randomly, or according to given mathematical functions, leading to infinite different networks. The methodology of the work is based on a combination of Object Oriented Programming and CAD software. That allows the design to be produced automatically within a few minutes, featuring a complex and irregular network of channels within any shaped implant, and thereby removes the need to spend innumerable hours attempting to manually model the same architectures. The output data should consist of a standard CAD file of reasonable size that can easily be converted into an STL format for manufacturing by LM.2. MethodologyDefining a network of channels involves the use of many geometrical parameters including: cross-sections of each channel, its number, size, shape and 3D path. Different algorithms and functions are needed to design these geometries.

Different algorithms and functions are needed to design these geometries. Those could be routines to perform different tasks such as generating parametric 2D and 3D complex curves, controlling their tangents and curvature, creating points by coordinates and on surfaces, offset and tangent planes, solids and holes of variable cross section, sketches, 2D and 3D constraints, or other functions used to set their location in the space. In order to test the feasibility of automated design by CAD, several different case studies have been investigated. These examples exhibit a progressive increase in geometry variables and a subsequent complexity related to an increasing number of design functions and routines in accordance with the input data set by the user. The efficiency of such routines is assessed by the output data including computational time and resources, and the resultant file size.

The efficiency of such routines is assessed by the output data including computational time and resources, and the resultant file size. The examples consisted of a small cube (10x10x10 mm) comprised of differently oriented channels of various shapes and sizes. There was no particular requirement to use a cubic shape - any other reference geometry could have been used, but this example provided a regular volume to direct attention onto other parameters, aside from the location of each channel. All the structures reported in this work have a regular location and therefore the position of all channels is fixed. That means that the networks obtained here are irregular in terms of geometry, but not in terms of position. The location of each channel constitutes a parameter which, like all other parameters, can be changed either randomly or according to any mathematical function.

The location of each channel constitutes a parameter which, like all other parameters, can be changed either randomly or according to any mathematical function. However, that was not considered in this section of work since it was not required to assess the proposed methodology. In a future work, the parameter will be changed in different ways, in order to define irregular distributions.In this work, all input parameters have been changed randomly, each within a user defined range, in order to increase the irregularity of the geometry and to simulate foam structures obtained by scaffolds made using conventional manufacturing techniques. All the structures shown here were designed using the CAD software Catia® V5 in conjunction with our programme routines and a standard Intel® Pentium® M processor 1.60 MHz laptop, with 752 MB of RAM.

All the structures shown here were designed using the CAD software Catia® V5 in conjunction with our programme routines and a standard Intel® Pentium® M processor 1.60 MHz laptop, with 752 MB of RAM. A standard computer has been used in this work to pursue novel automated design of TE scaffolds using standard equipment, without the need for high performance computers. The scaffolds consist of a group of open channels. In order to aid visualisation and understanding of each network, their complementary shape, formed by a network of rods, was also shown. These structures aim to represent the theoretical shape of the regenerated tissue, assuming no degradation of the scaffold.2.1 Evaluation of Pore SizeIn order to simulate an irregular shape in the cross section of each channel, closed 2D splines, consisting of 4 control points, were used.

Each control point was positioned as a vertex of a square and, in the first approximation, an estimation of all pore sizes was made, based on the side of each square and its area. The actual porosity is greater than this value, and can be computed using the mathematical formulation of Catia's spline curves. However, throughout this work the pore size was estimated using the above mentioned technique, because a more precise computation does not affect the design methodology itself. Fig. 4 gives an example of how the algorithm for generating such 2D splines works. The estimated pore size is defined by the highlighted square, while the actual size is the area enclosed by the spline. It is, of course, possible to define any other design layout--for example by using a different number of control points in which the coordinates do not necessarily form the vertex of a square, but can be changed randomly in different ranges.

It is also possible to control the wall thickness between two consecutive holes either by setting the minimum and maximum, and changing the value as an active constraint on a 2D sketch, or, by introducing other parameters such as the curvature on each control point, accessible in Catia's Application Programming Interface (API). The presented layout was chosen for its simplicity.STEPS TO CONTROL PORE SIZE AND SHAPE:Set pore size range (example 0.2 to 0.8 mm)Generation of a random number within this range (0.15 mm in this example)Creation of 5 control points (1st and 5th coincidents and equal tangent) positioned as 4 vertices of an ideal squareTo obtain different shapes, the tangents are alteredFig. 4 Steps for creating different porosity and shape in the 2D cross sections of each channel.3.

4 Steps for creating different porosity and shape in the 2D cross sections of each channel.3. Results3.1 Case 1This case presents a 10 mm cube with a regular distribution of holes of different shapes, obtained by randomly changing the value of the tangents at the control points as parameters. The pore-size range was set from 0.2 mm to 0.8 mm, but the actual pore size is slightly greater, since it was estimated by taking into consideration the square area formed by the 4 control points, which define the profile of each cross section. The pore size range chosen is suitable for regenerating bony tissue (Hulbert et al., 1970). The scaffold is illustrated in Fig. 5 (a), and its negative geometry in Fig. 5 (b). The details of this case are reported in Table 1.(a) (b)Fig.

The scaffold is illustrated in Fig. 5 (a), and its negative geometry in Fig. 5 (b). The details of this case are reported in Table 1.(a) (b)Fig. 5 Case 1: (a) scaffold, (b) theoretical regenerated bone.ConstantsVariablesInput Parameters: Cube side Pore size Location of holes Number of holes Tangents at 4 control points Output Parameters: Computational Time Porosity File size STL sizeValue 10 mm Range set from 0.2 to 0.8 mm Regularly fixed 5 Random Value 13 mins 55.27% 10.7 MB 9.95 MBTable 1 Case 1: main design and computational data.3.2 Case 2This case presents a 10 mm cube with a regular distribution of channels, defined by closed splines and with a step gradient in the cross sections. In each direction and for each hole, the shape changes randomly every 2 mm, by randomly altering the tangents at the 4 control points. The scaffold and its negative shape are illustrated in Fig.

In each direction and for each hole, the shape changes randomly every 2 mm, by randomly altering the tangents at the 4 control points. The scaffold and its negative shape are illustrated in Fig. 6 (a) and (b) respectively.Pore size range was set from 0.2 mm to 0.8 mm, considering the approximation of the square area formed by the 4 control points. Fig. 7 shows the step changes in the cross sections in more detail. The details of this case are reported in Table 2.(a) (b)Fig. 6 Case 2: (a) scaffold, (b) theoretical regenerated bone.(a) (b)Fig.

The details of this case are reported in Table 2.(a) (b)Fig. 6 Case 2: (a) scaffold, (b) theoretical regenerated bone.(a) (b)Fig. 7 Case 2: More details of the step changes of the cross sections for: (a) scaffold, (b) theoretical regenerated bone.ConstantsVariablesInput Parameters: Cube side Pore size Location of holes Number of holes Step distance Tangents at 4 control points Output Parameters: Computational Time Porosity File size STL sizeValue 10 mm Range set from 0.2 to 0.8 mm Regularly fixed 5 2 mm Random x 5 sections Value 15 mins 49.95% 15 MB 16.9 MBTable 2 Case 2: main design and computational data.3.3 Case 3This case shows a 10 mm cube with a regular distribution of straight and non-orthogonal channels with 2 cross sections, formed by 2D closed curves (splines with 4 control points). The linear path of each channel is defined by randomly connecting 2 different central points, located on opposite faces.

The linear path of each channel is defined by randomly connecting 2 different central points, located on opposite faces. The shape changes randomly by controlling the tangents at the control points. The scaffold and its negative shape are illustrated in Figs 8 (a) and (b) respectively. Pore size range was set from 0.2 mm to 0.8 mm, with a slightly greater actual size. Fig. 9 shows the linear channels in greater detail.(a) (b)Fig. 8 Case 3: (a) scaffold, (b) theoretical regenerated bone.Fig.

Fig. 9 shows the linear channels in greater detail.(a) (b)Fig. 8 Case 3: (a) scaffold, (b) theoretical regenerated bone.Fig. 9 Case 3: More detail of some linear channels and changes in the cross sections.The details of this case are reported in Table 3.ConstantsVariablesInput Parameters: Cube side Pore size Location of holes Number of holesLinear path for each channelTangents at 4 control points Output Parameters: Computational Time Porosity File size STL sizeValue 10 mm Range set from 0.2 to 0.8 mm Regularly fixed 7 Random association of the central points of 2 cross sections in opposite planes Random x 2 sections Value 20 mins 44.16% 35.7 MB 17.4 MBTable 3 Case 3: main design and computational data.3.4 Case 4This case consists of a network of channels made from two cross sections that are both defined by 2D closed splines. Tangents at the four control points are controlled randomly and a 3D spline defines a curved path.

Tangents at the four control points are controlled randomly and a 3D spline defines a curved path. As seen previously, there are many different ways to define a pattern and those presented in this and in subsequent cases are simplified solutions since any other arrangement could be set. In this example, all channels were set to converge at the geometrical centre of the cube, in order to improve the porosity and pore interconnectivity of the structure at the scaffold centre, thus mimicking biological internal bone structures (Fig. 10). This design could improve the flow of oxygen, cells and nutrients inside the scaffold and therefore influence the penetration depth of the new tissue, which can be problematic in linear foam scaffold structures.Fig. 10 Example of Metatarsal head µCT scan, illustrating internal bone structure.The curved path of each channel is defined by randomly connecting two different 2D curves at the same control points, with a 3D curve through the cube centre.

During this process, it ispossible that errors such as twisted and sharp geometries or cusps could cause the end of the programme and interrupt the automated design. To combat this, a routine was specifically designed to detect runtime errors and correct the geometry. This routine automatically detected any error in the geometry, and altered some input parameters until it produced an error-free geometry, without any manual intervention. The user can dictate which parameters are to be changed by such a routine. The tangents at the control points were set by default as was, subsequently, the pore size. That procedure increased computational time because of its need to calculate many different combinations of parameters in order to find a suitable one. Its advantage, however, is that it allowed the automated design of more complex channels involving curvature and 3D paths, along with multi-section elements.

Its advantage, however, is that it allowed the automated design of more complex channels involving curvature and 3D paths, along with multi-section elements. The scaffold and its negative shape are illustrated in Figs 11 (a) and (b) respectively. Again, the pore size range was set from 0.2 mm to 0.8 mm, with a slightly greater actual size. Fig. 12 shows some curved paths and their convergence to the scaffold centre in more detail.(a) (b)Fig. 11 Case 4: (a) scaffold, (b) theoretical regenerated bone.Fig.

Fig. 12 shows some curved paths and their convergence to the scaffold centre in more detail.(a) (b)Fig. 11 Case 4: (a) scaffold, (b) theoretical regenerated bone.Fig. 12 Case 4: More detail of some curved paths and their convergence to the scaffold centre.The details of this case are reported in Table 4.ConstantsVariablesInput Parameters: Cube side Pore size Location of holes Number of holesCurved path for each channelTangents at 4 control points Output Parameters: Computational Time Porosity File size STL sizeValue 10 mm Range set from 0.2 to 0.8 mm Regularly fixed 7 Random association of the central points of 2 cross sections in opposite planes, connected by a 3D curve through the centre Random x 2 sections Value 20 mins 33.56% 20.6 MB 13.6 MBTable 4 Case 4: main design and computational data.3.5 Case 5This case is similar to case 4, although the network is arranged differently.

In this example, the channels no longer converge towards the centre, and instead, their path is set randomly in a predefined range. The 3D spline that defines a singular channel is made of three points, the first and third of which are in the two cross sections, and correspond to two control points. The coordinates of the middle point are randomly variable within a range, dictated by user input. Here we have defined a range of ±20% the length of the cube, which means ± 2 mm in each direction in respect to the cube centre. However, any other value could be imposed. As in the previous case, it was necessary to include a routine to detect errors and automatically correct the geometry. The parameters modified by the routine were the tangents at the control points of the 2D splines and the pore size. The scaffold and its negative shape are illustrated in Figs. 13 (a) and (b) respectively. Pore size range was set from 0.2 mm to 0.8 mm, but again the actual size is slightly greater. Fig.

The scaffold and its negative shape are illustrated in Figs. 13 (a) and (b) respectively. Pore size range was set from 0.2 mm to 0.8 mm, but again the actual size is slightly greater. Fig. 14 shows some curved paths and their random connections in more detail.(a) (b)Fig. 13 Case 5: (a) scaffold, (b) theoretical regenerated bone.Fig.

Fig. 14 shows some curved paths and their random connections in more detail.(a) (b)Fig. 13 Case 5: (a) scaffold, (b) theoretical regenerated bone.Fig. 14 Case 5: More detail of some curved paths and their random connections.The details of this case are reported in Table 5.ConstantsInput Parameters: Cube side Pore size Location of holes Number of holes Control point of the 2 cross sections to be connected by the 3D curve Ideal Central VolumeValue 10 mm Range set from 0.2 to 0.8 mm Regularly fixed 7FixedVariablesCurved path for each channel±20% the length of the cube in each direction Random association of the central points of 2 cross sections in opposite planes, connected by a 3D curve through a random point within the Ideal Central VolumeTangents at 4 control points Output Parameters: Computational Time Porosity File size STL sizeRandom x 2 sectionsValue 22 mins 36.77% 35.5 MB 24 MBTable 5 Case 5: main design and computational data.4.

DiscussionThe experimental cases have shown that it is possible to achieve the automated design of regular distributions of irregularly and randomly connected channels using CAD software that supports full access to their modelling commands, through API functions such as CatiaV5.When compared to manual design, the advantages in terms of time are evident. The presented cubes are relatively small, but they could be extended to correspond to the size of a generic implant. The bigger the scaffold, the greater the time that can be saved by using automated design as compared to manual design. Fig. 15 shows the Case 4 scaffold extended to the size of 51 x 41 x 9 mm in order to fit the size of a craniofacial implant. It comprises a regular distribution of 420 channels in the XY plane, with each channel having two cross sections.

It comprises a regular distribution of 420 channels in the XY plane, with each channel having two cross sections. Trying to manually model such a network would be an extremely time consuming task, since for each channel, the user would need to find random combinations of parameters that do not lead to cusps, twists and other errors in the geometry, which may occur frequently in such features. In this case, the design time required by this methodology was 30 minutes, resulting in a CAD file size of 78.8 MB.Fig. 15 Example of Case 4 extended to fit the size of an external implant.However, the advantages of the methodology proposed in this work do not consist purely in time savings, although those are extremely significant.

This process could also permit a study of the influence of specific geometric parameters on tissue regeneration—such as the shape of channels and curvature by parametric studies, should all the geometries be fully controllable and reproducible by a layer manufacturing system. The values of all the random parameters can easily be stored in external files and used, for example, as input to redesign similar structures. Furthermore, instead of a random change, any mathematical function can be set for all or some of the inputs, in order to mimic the different densities found in bone structures. The possibilities to further develop this methodology are infinite and overcome the lack of flexibility and the rigid schemes generally encountered in CAD software. The flexibility and versatility of this methodology permits easy adaptation and modification, seen, for example, when the requirements for scaffolds are better defined.5.

The flexibility and versatility of this methodology permits easy adaptation and modification, seen, for example, when the requirements for scaffolds are better defined.5. ConclusionsSome factors are not incorporated into this methodology. The geometrical design features of the progressive case studies focus on the defined biological requirements identified in other works. It does not fully consider the computation of the mechanical properties of scaffolds, as this was not the purpose of this study. There is also a need for a more precise computation of the pore size and the definition of a parameter to measure pore interconnectivity. The level ofporosity is not yet set as an input parameter, but it is a consequence of the automated design process itself. To adopt that as a direct input, further investigations on the influence of the present input parameters on its resultant value are required.

To adopt that as a direct input, further investigations on the influence of the present input parameters on its resultant value are required. Ongoing work is being conducted on an efficient and optimised methodology for merging such networks of channels with a generic anatomical geometry from CT scan data, avoiding any Boolean operations. In addition, another class of networks, based on the irregular distribution of irregular channels is being investigated. In these networks, the location of each pore is considered as a random parameter and the connections between two cross sections are no longer located on opposite faces of the cube, resulting in random connections of the six faces. The final distribution of the channels within the cube will therefore be irregular.

tomated Fabrication of Complex Molded Parts Using Mold SOMA.G. Cooperl , S. Kang l , J.W. Kietzman l , F.B. Prinz l , J.L. Lombardi2 and L. Weiss31. Stanford University, Stanford CA 2. Advanced Ceramics Research, Tucson AZ 3. Carnegie Mellon University, Pittsburgh PAABSTRACTMold Shape Deposition Manufacturing (Mold SDM) is a Solid Freeform Fabrication technique for producing complex shaped fugitive wax molds. A vari~ty of castable. polymer and ceramic materials have been used to make parts from these molds. ThIS paper descnbes the Mold SDM method and an automated mold building machine based on a commercial CNC mill. Process steps, material selection and equipment issues are explained.

ThIS paper descnbes the Mold SDM method and an automated mold building machine based on a commercial CNC mill. Process steps, material selection and equipment issues are explained. Alumina, silicon nitride, polyurethane and epoxy parts with feature sizes ranging from 0.5 to 30 mm will be shown, as well as pre assembled mechanisms and multi-material parts.INTRODUCTIONCeramic materials have excellent high temperature properties, high strength and stiffness, low density, and good chemical resistance. These properties make ceramic materials very attractive for a wide range of engineering applications. However, two principal factors currently limit the use of structural ceramics: the low toughness of ceramic materials and the difficulty of making high quality complex shapes. The Mold SDM process addresses the latter issues.Traditional ceramic manufacturing processes, such as machining of green ceramic blanks, work well for making simple shapes.

The Mold SDM process addresses the latter issues.Traditional ceramic manufacturing processes, such as machining of green ceramic blanks, work well for making simple shapes. Solid FreeformFabrication (SFF) processes, based on layered manufacturing, can build up complex shapes but, in general, produce parts with poor surface finish, often displaying a stairstep effect on surfaces that are not vertical or horizontal. The layer boundaries are also potential ~ources of.defects thatcan reduce the mechanical properties. Mold SDM was developed to address these issues. MoldSDM is a layered manufacturing process, so it can build complex shaped parts. ·However, it·isalsoanadditive-subtractiveprocesswhich uses CNC milling to accuratel)' shape all surfaces as they are built to minimize thestairstep effect.

·However, it·isalsoanadditive-subtractiveprocesswhich uses CNC milling to accuratel)' shape all surfaces as they are built to minimize thestairstep effect. Although the mold is built in layers, there will be no layer boundaries in the finished part since the final part is cast monolithically.In addition to ceramic parts, Mold SDM can be used to make parts from a variety ofcastable materials including polymers such as polyurethane and epoxy.. . This paper describes two aspects of the MoldSDM process. First is the Mold SDM process Itself In terms of process steps and the related materials and process issues. Second is the automation of the process and the construction ofan automated machine that implements the mold building phase of the process.THE MOLD SDM METHODMold SDM is a variation on Shape Deposition Manufacturing (SDM). SDM is an additive subtractive layered manufacturing process which has been used to build a variety of metal and polymer parts [1,2].

SDM is an additive subtractive layered manufacturing process which has been used to build a variety of metal and polymer parts [1,2]. Most SFFtechniques decompose the modeLinto relatively thin planarlayers of uniform thickness. In SDM. however, the layers are three-dimensional, may be of arbitrary721thickness and do not needlo be planar.Thj~a(iaptiYe decomposition allows th~ ~umberof layers ~o be minimized which in turn leads toreduc~dbuildtimes. In Mold SDM fugItIve molds are bUIlt using SDM techniques and these are then used to cast a variety of part materials.Figllre 1illllstrates. an.ex.a1I.lPI¢Qftbe~91c:tSJ.i)~bllj1djJ).g.~e<Illepcefpr.a ••siInple.partwith three layers. A mold is constructed from mpldmaterialsllITounding temporary part material which defines the mold cavity. Instepsl thr0tlghi~themoldis built llP layer by layer. Removal of the temporary part material, in step 9, le.aves a moldreadyfor casting in step 10.

Instepsl thr0tlghi~themoldis built llP layer by layer. Removal of the temporary part material, in step 9, le.aves a moldreadyfor casting in step 10. Oncethe part material has cured there are a range ofproce~sing.optjollS available. The figure shows two possible alterpatives.• In the first option (lla and 12a), the mold material is removed and then finishing operations, such as removal of runners and gates, are performed.• In the second option (lIb and 12b),the mold material is used as fixturingduringfinishing operations and is removed afterwards.1234512867891011b12bMold MaterialTemporary Part Material_Part MaterialFigure 1: Example Mold 80M process sequenceThe main advantage of Mold SDMover SDMand other layered processes is that since the final part is cast ffi<.:molithically there will be no layer boundaries in the finished part.

This is particularly advalltageous for flaw sensitive materials, such as ceramics, orfor materials with poor interlayerbonding. In comparison with direct machining of ceramic green materials, Mold SDM also minimizes the amount of part machining required which is an advantage for materials that are difficult or expensive to machine, such as ceramic green materials.Mold SDM has several limitations overSDM. First, a third compatible material,· the temporary part material, is required. The additional materials compatibility and processing requirements may restrict the range of materials that can be used. Second,tb.ere are the extra casting and mold removal steps Which. will increase processing time.

Second,tb.ere are the extra casting and mold removal steps Which. will increase processing time. Finally, mold fillingissues may.limit part geometry, although in many cases sprues and vents canbe added to ensure complete mold filling.Materials IssuesThe capability to make quality molds from which quality parts can be produced depends on a number of mold and temporary materialproperties [3]. The materials must have low shri~age to reduce warping and distortion. They must bond to each other so that they do not delaminate and are able to withstand cutting forces. They must machine well so that good surfaces can be produced. They must be chemically compatible with each other so that they do not damage each other, both during the build phase and the two etching steps.In .addition to these basic requirements there are a number of other material properties which are desired to make the process more rapid or efficient.

Being able to etch the temporary partmaterial and later the mold material quickly and easily saves time. Low cost, non-toxic materials make the process more economical.The main issue with the mold material is the tradeoff betwe~n shrinkage and machinabil~ty. In general machinable waxes have high shrinkage while low shrmkage waxes tend to machme poorly. Table 1 shows a range of waxes that have been tested for use in Mold SDM.WaxKindt-Collins Master Protowax Kindt-Collins Master File-a-wax Freeman machinable 25% Kindt-Collins. Master File-a-wax 75% Kindt-Collins Master ProtowaxMelting Point toe) 60 105 105 80Shrinkage(linear %) 0.46 2.60 2.91 0.80Table 1: Waxes tested for use in Mold SDMCurrently the preferred wax is the 25/75 mix of Kindt-Collins Master File-a-wax and Protowax. The File-a-wax is a machinable wax, the Protowax is a casting wax.

The File-a-wax is a machinable wax, the Protowax is a casting wax. By mixing the two it is possible to obtain a better tradeoff between machinability and shrinkage than is available in a commercially available wax.Melting point and melt viscosity are two other important properties. A lower melting point wax is desired because this will reduce the heat input into the previous layers during wax casting thus reducing mold warpage. However, the wax must have a high enough melting point to be able to withstand the cure conditions for the final part material. Many materials require an elevated temperature cure or have an exothermic curing reaction. Low melt viscosity is beneficial since this will allow the wax to more easily fill fine features. It will also make wax removal by melting easier as the wax will run off the part more readily.The temporary part material must be able to accurately define the mold cavity.

It will also make wax removal by melting easier as the wax will run off the part more readily.The temporary part material must be able to accurately define the mold cavity. To do this the temporary part material must be machinable so that fine features can be made and heat resistant so that it can withstand wax deposition conditions. Similar to the case for wax formulations, it was found that a better material could be obtained by mixing commercially available soldermask formulations. Mixing brittle and rubbery varieties produced a tougher, more machinable soldermask. One drawback of using this mixture, however, is that the viscosity gradually increases over time. This mixture also suffers from low heat resistance and low cure depth. Low heat resistance means that features are more likely to be distorted when hot wax is cast over the soldermask. Low cure depth reduces build rate because thick layers must be built up in several steps.

Low cure depth reduces build rate because thick layers must be built up in several steps. The current formulation can be cured up to 1.5 mm deep in 2 minutes.Process IssuesIn Mold SDM wax layers are deposited by mass casting. This enables much more rapid material deposition than is possible with bead extrusion systems. In order to prevent the wax from flowing away from the desired deposition area some sort of containment must be used. In Mold SDM this is done by building up thin wax walls around the deposition area. These walls are built up by extruding a higher melting point wax that has sufficient viscosity to form a narrow bead. The wax is deposited at a temperature only slightly above the melting point so that it will solidify rapidly and form precise walls.

The wax is deposited at a temperature only slightly above the melting point so that it will solidify rapidly and form precise walls. Once the walls have been formed, the lower melting point mold wax is mass cast inside the walled off areas and allowed to cool.723In orderto reduce thermal stresses and minimize the potential of damage to previously built layersit is desirable to cast the wax la,yers .at as Iowa telllperature as possiple. However at lower temperatures there will be less remelting of the previous layer, therefore the interlayeradhesion will notbe .asggod..An .9gtion is.touse.prehe~til1giofthepr~vig~s.layer,.using anIR lampforexalllple, t?iwaxl11u.p·.·the/surfa.97is?t~~tit~illJ~l1171tlllor~9a~ily•.•and r~~1.lce.the••Jeqlliredd~position temperatures. For example th~ Kindt~Cgl1insMaster Fi1e"a~waxdepositiontemperature can be reduced from 180°(2 to 160°C. Onedisadvantag~ of pr~heatin.g is that interlayer discontinuities tend to be greater.

For example th~ Kindt~Cgl1insMaster Fi1e"a~waxdepositiontemperature can be reduced from 180°(2 to 160°C. Onedisadvantag~ of pr~heatin.g is that interlayer discontinuities tend to be greater. This is believedto be dueto soft~ning of the surface layer that is then distorted as the new wax layer cools and shrinks above it.The cooling time between wax deposition and machining accoun.ts for a significant fraction of thetotal part build time. Wax layers must usually be allowed to cooHor 2 hours for a typical 5 10 rom thick layer. Thicker layers.will take longer.. In.order to reduce this time and increase the process rate, alternative ways of speeding the cooling ofthe wax are currently being investigated. By simply blowing cool air over the surface it is possible to reduce cooling times by a factor of 2 4.PROCESS AUTOMATIONTo improve the rate of production .and the quality of parts made by Mold 8DM, an automated mold-making machine has been built.

The machine, shown in Figure 2, is based on a commercially available Haas VF-OE 3-axis CNC milling machine. The mill was modified by the addition of mat~rialdeposition and curing hardware. Ultimately one could produce either complete Mold 8DM machines or retrofit kits that could be installed on existingCNC machines.Figure 2: Automated Mold SDM machine picture(Figure 3 shows .. the main components of the Mold 8DM machine. The CNC mill's machining functionality is used to perform all shaping .operations. The add~on equipment together with the mill's positioning capabilities performs material deposition and curing.The CNC mill is used to perform all XY motion, as well as the Z motion for the dispensers. The vertical positionin.g of the dispensers is critical since these need to be able to track the part surface accurately in order to deposit material in the right locations.

The vertical positionin.g of the dispensers is critical since these need to be able to track the part surface accurately in order to deposit material in the right locations. By attaching the dispensers to the ram of the mill, one can use the Z positioning capabilities of the machine. This724makes implementation much easier since it is not necessary to build additional Z P?siti0!1i~g hardware that is coordinated with the XY motion of the. mill. The ram does have a weIght hffilt, however, so the hardware attached to it should be as lightas possible.Soldermask/Tank"Dispensers (2 wax, 1 soldermask)Figure 3: Automated Mold SDMmachine schematicThe dispensers are attached to the ram via pneumatically actuated linear slides to produce an up and a down position. The dispensers are normally held in the up position to keep them clear of the part during machining. To deposit material, the appropriate dispenser is moved into the down position and material deposition begins;.

To deposit material, the appropriate dispenser is moved into the down position and material deposition begins;. When finished the dispenser is retracted back to the up position.The lights do not need to follow the surface. of the part and so their Z pmdtioning requirements are less strict. They are usually positioned ata set height above the surface and turned on while the part moves back and forth undetrleath'forthese reasons,andto minimize the amount ofequipment mounted on the ram, a separate linear actuatormoves the lights up and down.The Slautterback KB 10 wax dispensers are commercially available hot melt dispense units which have a heated melt unit and a pump which forces the molten wax along a heated hose to a dispense valve. Since two types of wax mustbe deposited, one for building containmentwalls for casting and one for building the molditself, two wax dispensers are required.The soldermask dispenser consists of a pneumatically operated EFD 725DA-SS dispensevalve.

Soldermask is supplied from a pressurized reservoir.The Uvexs OCU-12C ultraviolet light is used to cure the soldermask. It is a line source 12 inches long with a power intensity of 300 W/in. The light contains anIR filter tominitnize the amount of IR light emitted. It is important to minimize the amount of IRlight em.ittedbecause this will heat the previous layers and may cause warping, or even melting inextrem~cases.The Research Inc. 4453-A-IO-06 IR area heater is us~dto preheat the substrate before deposition of wax. This is anIR heat lamp with six 1000WbulbsthatcoversalOjn.by6~p.,area.Si~ye~either light covers the entire 12 in. by 12 in. area oLa pallet, curing or preheating operations are performed by scanning the part back and forth under the light.725Control of the integrated CNC deposition/milling machine is performed using both the CNC mill's built-in controller and an external computer, as shown in Figure 4.

The computer controls the deposition and curing apparatus via digital and analog 110 and sends motion commands to the CNC controller by downloading CNC program files via an RS232 serial line. The mill also has four relay closure outputs and a digital input that can be controlled and monitored from within the CNC program in order to synchronize the mill's motion with the deposition and curing operations. During a curing operation, for example, the mill will move the part below the UV light, set one of its outputs and wait. When the computer detects that output signal, it turns on the UV light and signals the mill to continue. The mill then begins to move the part back and forth under the light.ComputerRS232 SerialCNC MillI/O ElectronicsRelay ClosureDigital and/or AnalogDeposition Equipment and Li htsFigure 4: Machine control schematicEXAMPLE PARTSval'iety. of castable part matel'ialshav~.lJeenusedto make parts llsingMoldSDM. PolYITIel' parts.

of castable part matel'ialshav~.lJeenusedto make parts llsingMoldSDM. PolYITIel' parts. have been madel:lsingp()lyur~thanesand.epoxy. Ceramic parts have been made usingal14mina and silicon nitridy gelcllstingformulations. Pre~assembled mechanisms have also been built in both polymer and ceramic materials.Figure 5 shows three multi-material turbine assemblies. The epoxy rotor consists of eight curved blades between two discs. A captive polyurethane shaft passes through ahole in the center of the rotor. The rotor is free to spinaro.undtheshaft. The diameter ofthe rotors is 30 mm. These assemblies were built with radiaIc!earances as small as 0.20 mm between the shaft and rotor.In an effort todeterUlineminimumfellture siles for MoldSDM a set of miniature alumina turbines were built. These are shown in Figure 6 infrollt of a full size part. The miniature turbines are 7 mm in diameter and 3 mm tall.

These are shown in Figure 6 infrollt of a full size part. The miniature turbines are 7 mm in diameter and 3 mm tall. The section thickness is 0.5 mm.Figure 7·shows a simplified aircraft engine vane part made of alumina. This part is about60mm tall.Fig14re.8 shows a part called theInchworm. It featur~s two pairs of wheels with ratchetsmounted toa flexible backbone. The ratch~ts prevent the wheels from turning backwards. Pressing down on the backbone rolls the front wheels·forward, releasing the backbone allows the rear wheels to roll forward in turn, thus moving the Inchworm forward. Unfortunately the wrong clearances were used in the actual parts and as a result the ratchets do not work. The Inchworms726were also the first Mold SDM parts made using multiple cavity molds. This allowed the wheels to be cast using colored polyurethane to distinguish them from the backbone.The part shown in Figure 9 is a pitch shaft for a missile guidance system.

This allowed the wheels to be cast using colored polyurethane to distinguish them from the backbone.The part shown in Figure 9 is a pitch shaft for a missile guidance system. This partdemonstrates the ability of Mold SDM to build curved surfaces without stairsteps.Figure 5: Multi-material turbinesFigure 6: Miniature alumina turbinesFigure 7: Alumina vaneFigure 8: Multicolor polyurethane InchwormFigure 9: Alumina pitch shaft727CONCLUSIONThe Mold SDM process was developed to enable the fabrication of high quality complex ceramic parts. It is an additive-subtractive layered manufacturing process that builds complex fugitive molds. The use of CNC milling to shape all geometry results in smooth accurate surfaces. Since the final part material is cast monolithically there will be no layer boundaries in the finished part.The Mold SDM process has been automated by the addition of deposition and curing hardware to a commercially available CNC milling machine.

Parts have been made without manual intervention using this machine.Alumina and silicon nitride ceramic parts have been made by gelcasting. Other castablematerials, such as polyurethane and epoxy, have also been used successfully.ACKNOWLEDGMENTSThe authors would like to acknowledge Shelley Cheng for building the miniature turbine parts as well as David Miller and Brad Levin for building the Inchworm parts. The authors would also like to acknowledge Tom Hasler for his advice and help with all aspects of machining and CNC mill operation. The authors would also like to acknowledge Larry Schultz, Blake Tennison, and Mark Carter for their assistance with process automation equipment.This work was supported by the Defense Advanced Research Projects Agency under contract NOOOI4-96-1-0625. Additional support was provided by a Department of Energy Integrated Manufacturing Pre-Doctoral Fellowship.REFERENCES[1] R. Merz, F. B. Prinz, K. Ramaswami, M. Terk, and L. E.

Additional support was provided by a Department of Energy Integrated Manufacturing Pre-Doctoral Fellowship.REFERENCES[1] R. Merz, F. B. Prinz, K. Ramaswami, M. Terk, and L. E. Weiss, "Shape Deposition Manufacturing," Proceedings ofthe 1994 Solid Freeform Fabrication Symposium, The University of Texas at Austin, August 1994, pp. 1-8.[2] F. Prinz and L. Weiss, "Method for Fabrication of Three-Dimensional Articles," U. S. Pat. No. 5,301,415, April 12, 1994.[3] J. W. Kietzman, A. G. Cooper, L. E. Weiss, L. Schultz, 1. L. Lombardi, and F. B. Prinz, "Layered Manufacturing Material Issues for SDM of Polymers and Ceramics", Proceedings ofthe 1997 Solid Freeform Fabrication Symposium, The University of Texas at Austin, Austin, Texas, August 1997, pp. 133-140.728.

tomated Fabrication of Monolithic and Ceramic Matrix Composites via Laminated Object Manufacturing (LaM)Don Klosterman, Richard Chartoff Nora Osborne, George Graves Rapid Prototype Development Laboratory and Ohio Rapid Prototype Process Development Consortium University of Dayton 300 College Park Dayton,OH 45469-0130AbstractThis report summarizes recent developments in a research program for fabricatingadvanced monolithic and ceramic matrix composite parts using Laminated Object Manufacturing (LOM). Both silicon carbide (SiC) and SiC/SiC composites are discussed. The LOM process is used to produce green forms that are then densified using various post processing operations. The monolithic ceramic LOM process was advanced through the implementation of an automated solvent spray bonding step, significant improvement in decubing with new software, and an intensive round of mechanical characterization. The LOM process for making CMC green forms is fully developed.

The LOM process for making CMC green forms is fully developed. This entailed implementing a process for making suitable SiC fiber preforms, a laser cutting capability, a decubing strategy, and a binder resin cure procedure. Further research is ongoing for the post processing pyrolysis and reaction bonding steps as discussed herein.IntroductionThe rapid prototyping and manufacturing of monolithic ceramics and ceramicmatrix composites (CMCs) is being studied by a university/industry collaborative team lead by the University of Dayton and Helisys, Inc. This technology, based on Laminated Object Manufacturing (LOM), will be capable of producing high performance, near-net shape structural ceramic parts. The standard LOM process [1] is illustrated in Figure 1.

The standard LOM process [1] is illustrated in Figure 1. The technical objectives of the program are: 1) to produce CMC parts with structural integrity similar to CMC parts made by current fabrication techniques, and 2) to develop a commercially viable machine capable of automated fabrication.

The rationale for using this process for composite fabrication is LOM's inherent capacity to handle flat sheet materials, to produce geometrically complex objects, and to operate with a high degree of automation, all without special tooling or fixturing.537The program involves a multidisciplinary effort that can be decomposed into thefollowing interrelated sub activities:Machine design and control Software modifications and enhancements • • Ceramic preform formulation and manufacture • Ceramic LOM process development • Laser cutting • Binder burnout • Ceramic precursor pyrolysis • Ceramic densification via reaction bonding • Mechanical testing and characterization • Part building and demonstrationPast efforts in all of these areas have been summarized in previous articles [2-6].

In this paper, recent progress in the areas of ceramic preform formulation and manufacture, ceramic LOM process development, ceramic precursor pyrolysis, and mechanical testing and characterization will be presented.Monolithic CeramicsAutomation of Solvent SprayThe LOM process for monolithic ceramics is being developed using SiC as afocus material system. Ceramic tapes of 250 f.1m thickness comprised of a bimodal SiC powder (3 f.1m and 60 f.1m), carbon and graphite powders, and a polymeric binder system are fabricated with a standard doctor blade tape casting process. As described in previous reports [3], the tapes are laminated on a standard LOM2030 using a solvent assist spray technique. A hobby/artist's airbrush is used to apply a fine spray of butanol to the resin rich side (bottom) of the ceramic tape.

A hobby/artist's airbrush is used to apply a fine spray of butanol to the resin rich side (bottom) of the ceramic tape. The solvent acts as a tackifier to enhance the strength of the lamination, lower the required lamination temperature, and drastically increase the reliability of repetitive layer lamination, all compared to lamination without using a solvent. Butanol as a solvent has several advantages: low vapor pressure so that it doesn't evaporate during the spraying operation, reduced toxicity compared to aromatic solvents, and reduced flammability compared to lighter solvents.A semi-automatic spraying process was developed to facilitate efficient, reliable, and reproducible part building (see Figure 2). The airbrush nozzle was mounted to the laser gantry of a LOMIOl5 machine. Solvent was fed to the nozzle through a flexible tube connected to a reservoir.

The airbrush nozzle was mounted to the laser gantry of a LOMIOl5 machine. Solvent was fed to the nozzle through a flexible tube connected to a reservoir. The spray stream, directed normal toward the platform, was manually turned on and off with a solenoid switch regulating the gas pressure. A standard LOM software routine (i.e., "Cut Tiles") was used to move the nozzle in a raster pattern while the solvent spray was activated. Tapes were placed and removed manually from the platform. The sprayed tapes were placed solvent side down on the part being built on an adjacent LOM2030.538The following parameters affect the solvent spray process: internal nozzle adjustments, gas pressUre, gantry speed, platform to nozzle distance, raster spacing, overall area sprayed. All these were determined by trial and error within a day's time, and they remain the same for all parts with the exception of the overall sprayed area.

All these were determined by trial and error within a day's time, and they remain the same for all parts with the exception of the overall sprayed area. This procedure has proven itself effective: parts of up to 50 layers have been built in an efficient, methodical, and reliable manner without any lamination flaws. Although this process is only semi-automatic, personnel with minimal training can easily handle it. Only one person is required to run the process.DecubingThe application of solvent to improve layer adhesion has important implications on the decubing step. The solvent assists in partially fusing the layers together, thus blurring the boundary. In general, the ease of decubing relies on the existence of an interface which, regardless of whether it is strong or weak, must be distinct.

In general, the ease of decubing relies on the existence of an interface which, regardless of whether it is strong or weak, must be distinct. After building several different ceramic parts with various geometries, it became apparent that the decubing step was difficult and tedious at best, and in many cases ceramic parts could not be decubed without severe damage. To address the problems of decubing ease and damage, two strategies were investigated: adaptive crosshatching and adaptive surface burnishing. Both these strategies are similar and required additional software to be developed. These strategies, illustrated in Figure 3, involve selectively modifying the region of overlap between cube area and part area in successive layers. In the case of adaptive crosshatching, the area is cut with a fine (0.5 mm) crosshatch. With adaptive surface burnishing, the surface of the overlap zone is burnished with the laser, removing the binder from the surface while leaving the powder.

With adaptive surface burnishing, the surface of the overlap zone is burnished with the laser, removing the binder from the surface while leaving the powder. The thin layers of loose powder serve as a release mechanism during decubing.Software developed by Helisys, Inc. for adaptive crosshatching was tested onthree parts shown in Figure 4. This strategy was largely successful. Each part was decubed with relative ease in ten or twenty minutes. More importantly, there were no flaws introduced from cubes sticking to the part or layers delaminating. Some of the success is directly attributable to the strength of the layer-to-Iayer bond as a result of the solvent lamination process. Upon closer examination ofthe surface with a microscope, it is evident that the pattern of fine crosshatches remains on the stair steps (see Figure 5). Thus, removal of this pattern and surface fInish become issues to consider.

Thus, removal of this pattern and surface fInish become issues to consider. Depending on the intended use of the part, it may not be necessary to improve the surface finish. However, if finishing is necessary, it is expected that normal grinding and polishing can be used to smooth the surface with no more difficulty than for parts made with other rapid prototype techniques. Therefore, the existence of the fine crosshatch pattern is not expected to pose any serious disadvantages.A preliminary analysis of adaptive surface burnishing using simple geometrieshas revealed that surface burnishing is at least as effective as adaptive crosshatching. The advantages of surface burnishing compared to adaptive crosshatching are 1) burnishing the overlap zone generates less dust compared to cutting; and 2) the part surface is not left with a fine crosshatch pattern.539Mechanical PropertiesAn intensive characterization of the mechanical behavior of LOM SiC parts wasmade.

MOR bars (4mm x 3 mm x 50 mm) were fabricated using the latest version ceramic LOM process and densified via reaction bonding. Over 50 bars were made in each of 6 different LOM runs. The results of the 4-point flexure tests are summarized in Table 1. It is important to note here that these specimens were tested as-received without any polishing.Table 1 : four-point flexure test results for as-received (unpolished), reaction bonded, monolithic SiC specimens made with LOM.LOM Run #1,2, & 3 4,5, & 6 1 2 3 4 5 6 5 6#Displace- specimens ment rate (cm/min) 0.5 0.5 0.5 0.5 0.5 0.05 0.05 0.05 0.005 0.00510 each run 10 each run 23 22 24 25 7 8 22 21Flexure Strength (MPa)158 152 157 151 158 142 142 154 152 165cr (MPa)34 25 22 22 18 23 20 10 21 23Weibull Modulus5.2 5.7 lOA 8.3 9.0 6.6 8.2 17.5 8.7 7.6The 4-point bend strength of the unpolished SiC specimens is about 150 MPa.

It is not valid to compare this strength to that of commercially available SiC because of the effect of polishing and chamfering in the case of the latter. In general, the strength of reaction bonded SiC is not expected to be as high as sintered SiC, which is approximately 400 MPa [7]. With proper preparation, the bend strength ofLOM SiC is expected to significantly increase from 150 MPa, most likely by a factor of two. If this is found to be the case, then the bend strength of reaction bonded SiC will compare favorably with that of similar commercially available material made with conventional techniques.The consistency of the bend strength from LOM run to run and within each LOM run is noteworthy. This consistency is characteristic of SiC as a material in general. The fact that the bend strength does not change with strain rate indicates that this material does not exhibit slow crack growth, another characteristic of SiC in general.

The fact that the bend strength does not change with strain rate indicates that this material does not exhibit slow crack growth, another characteristic of SiC in general. Thus, in two more instances, LOM SiC compares favorably with commercially available material.A photomicrograph of reaction bonded SiC (Figure 6) illustrates little evidence of a layering effect. In addition, low porosity indicates effective infiltration during reaction bonding.540Ceramic Matrix Composites (CMCs)The fabrication of continuous fiber CMCs with LOM is another objective of thisprogram. SiC/SiC is being used as a focus material system. From the onset, the two most critical issues were considered to be 1) the availability of preforms, suitable for LOM, which incorporate a high volume fraction of continuous SiC fibers; and 2) near net-shape densification of the CMCs.

Described in this section is the development of a suitable SiC/SiC preform for LOM and the various stages of the LOM process that dictated the preform's design. Issue 2 is being addressed through application of the reaction bonding process, as demonstrated with monolithic SiC, although detailed results are not yet available.The approach to fabricating CMCs with LOM involves the layup of separate,alternating monolithic ceramic tape layers and fiber/resin prepreg layers. The development of a SiC fiber (NicalonTM, Dow Corning) and furfural resin (FurCarbR UP-440, QO Chemicals) prepreg was described in a previous report [3]. The furfural is a thermoset resin that serves a dual role: as a binder during LOM fabrication and as carbon source for the reaction bonding process. Layers are bonded through heat and pressure applied by a flat heating plate or iron; no solvent is necessary.

Layers are bonded through heat and pressure applied by a flat heating plate or iron; no solvent is necessary. The fiber/tape layup is subsequently post-cured and pyrolyzed, followed by reaction bonding.Early attempts to fabricate CMCs with separate layers of tape and fiber werelimited by two practical problems, illustrated in Figure 7. First, the required laser power to cut through a fiber layer is much higher than for a monolithic tape layer. Although the laser power can be adjusted accordingly, it is not possible to handle the problems associated with random variations in the layer thickness or fiber content. Specifically, if the laser does not cut completely through a fiber layer, it will not be possible to simply "snap off' the adjacent waste cube during decubing due to the connected fibers.

Specifically, if the laser does not cut completely through a fiber layer, it will not be possible to simply "snap off' the adjacent waste cube during decubing due to the connected fibers. On the other hand, it the laser over compensates and delivers more power than is required to a fiber layer, the monolithic tape layer below will be severely damaged.The second problem associated with the use of separate layers is that the adhesionof a fiber layer to the top of a tape layer is significantly weaker than to the bottom of a tape layer. This effect is caused by the ceramic tape which contains a resin-rich side on the bottom. The decubing is impractical because it is difficult to separate layers at the strong interface without delaminating the adjacent weak interface in the process.

The decubing is impractical because it is difficult to separate layers at the strong interface without delaminating the adjacent weak interface in the process. Thus, the problems can be generalized as an inherent difficulty of working with two separate layers that have significantly different bonding and cutting characteristics.The situation is resolved by combining the separate tape and fiber preforms intoone preform during the prepreg B-staging cycle, as illustrated in Figure 8. This new preform, the "tape-preg", is a simple fiber-on-tape laminate that solves both the laser cutting and laminating problems by eliminating the weak interface so that each layer contains both fibers and monolithic tape. The laser cutting is improved because the fiber component is the first to be cut. Therefore, if insufficient power is delivered to cut the541entire tape-preg layer, the remaining material will be ceramic tape which is easily "snapped-off' during decubing.

Therefore, if insufficient power is delivered to cut the541entire tape-preg layer, the remaining material will be ceramic tape which is easily "snapped-off' during decubing. If too much laser power is delivered accidentally, the tape-preg layer immediately below will be minimally damaged, since its surface is comprised of fibers. The decubing is also improved because all the layer interfaces are of the same strength. The only practical complication is that the part layer thickness is necessarily doubled, unless thinner monolithic tapes and fiber prepregs are used. In this study, the tape-preg thickness was 0.5 mm (0.25 mm ceramic tape and 0.25 mm fiber prepreg).To test the effectiveness of the SiC tape-preg, a small scale turbine engine seal(Figure 9) was fabricated with the "Table-Top LOM" (TTLOM) system at University of Dayton [4]. This experimental system was assembled to study laser cutting of high performance fibers and build CMC parts throughout the research program.

This experimental system was assembled to study laser cutting of high performance fibers and build CMC parts throughout the research program. The apparatus combines a copper vapor laser with a computer controlled XY gantry, manually adjustable z-platform, and manual layer lamination via hot iron. With this apparatus, one is able to obtain high quality laser cuts, illustrated in Figure 10, as defined by steep cut channels, no fiber end damage, and a minimal resin burn-back zone. The turbine engine part was built with seven layers of SiC tape-preg alternating in a 0/90° arrangement. It was built and decubed without any problems. The decubing process was more similar to that for LOM paper parts rather than that for LOM monolithic ceramics. The reason for this is that the tape-preg layers maintain a distinct interface, Le. the monolithic layers do not migrate into the fiber layers and vice-versa.

The reason for this is that the tape-preg layers maintain a distinct interface, Le. the monolithic layers do not migrate into the fiber layers and vice-versa. For this reason, the use of adaptive crosshatching was not required.After decubing, SiC/SiC parts must undergo a resin cure cycle to complete thecure of the furfural resin and to provide additional consolidation. The turbine engine seal was fit with a negative made from LOM paper, placed in a vacuum bag set-up, and cured in an oven using the following cycle: 25-200°C for 7 hrs; 200°C for 1 hr. After this cycle, the part was physically robust. The microstructure is given in Figure 11. Notice that the layers are well compacted with little or no porosity present.The next step in the post-processing cycle was simultaneous resin pyrolysis andbinder burnout. This step was evaluated by pyrolysis of free-standing, 6-layer, I-inch SiC/SiC squares in flowing argon.

This step was evaluated by pyrolysis of free-standing, 6-layer, I-inch SiC/SiC squares in flowing argon. Early studies revealed that the laminates were prone to delamination during the pyrolysis cycle. A variety of cycle conditions were tested under the initial suspicion that the problem was related to inability of the pyrolysis products to successfully diffuse out. However, two general observations were made: the delaminations always occurred within the tape-preg layers, not between the tape-preg layers; and the layers tended to curl perpendicular to the fiber direction. Thus, it is more likely that the problem is caused by the shrinkage of the furfural resin and the differential bond strength throughout the CMC.

Thus, it is more likely that the problem is caused by the shrinkage of the furfural resin and the differential bond strength throughout the CMC. Although the bond within the tape preg layers is relatively strong during the LOM layup process, it becomes a relatively weak bond during the subsequent cure and pyrolysis cycle as the furfural resin cures evenly throughout the bulk.542The delamination problem worsens with increasing mass loss during the pyrolysis cycle. When the cycle is stopped at 250°C, there is 6% mass loss in the composite and it is possible to obtain an undamaged laminate using a slow heating cycle (2.5-5°C/hr). However, when the cycle is advanced to 300°C, there is 9.5% mass loss and the sample is delaminated. Based on thermogravimetric analysis, the laminates are expected to lose 20% mass through 700°C. It appears that to solve this problem a strategy involving only adjusting the heating rate will be unsuccessful. Alternative solutions are suggested.

It appears that to solve this problem a strategy involving only adjusting the heating rate will be unsuccessful. Alternative solutions are suggested. The first involves applying pressure to tape-preg laminates during pyrolysis. In an initial test, a sample that was placed under 15 psi pressure still delaminated during the pyrolysis cycle carried out to 300°C. Higher pressures will be investigated. A second strategy involves lamination of fiber layers only, omitting the tape layers. This strategy was successful in an initial study when a 10-layer laminate made of fiber layers survived the pyrolysis cycle through 550°C without delaminating. A third approach is to reconsider bonding separate layers of tapes and prepregs. The most appropriate strategy of these three is now under evaluation.ConclusionsThe monolithic LaM process was made more reliable and less labor intensivethrough recent hardware and software improvements.

Monolithic SiC parts can now be fabricated with ease via LaM and post processed to dense ceramic through reaction bonding. These parts exhibit comparable mechanical behavior to commercially available SiC specimens fabricated with traditional processes. The LaM process for making flat layer, continuous fiber SiC/SiC composites has been demonstrated to the point of making fully cured green parts. Additional work is needed to develop a fully effective pyrolysis cycle. Once achieved, reaction bonding of SiC/SiC is expected to be feasible using monolithic SiC technology.References1. Helisys, Inc., 24015 Garnier S1., Torrance, CA, 90505, USA.2. Klosterman, D., R. Chartoff, N. Osborne, G. Graves, A. Lightman, and G. Han, "Laminated Object Manufacturing (LaM) of Advanced Ceramics and Composites," Proceedings ofthe 7th International Conference on Rapid Prototyping, University of Dayton and Stanford University, San Francisco, CA, April 1997.3. Klosterman, D., R. Chartoff, N. Osborne, G.

Klosterman, D., R. Chartoff, N. Osborne, G. Graves, "Laminated Object Manufacturing, ANew Process for the Direct Manufacture of Monolithic Ceramics and Continuous Fiber CMCs," Proceedings ofthe 21st Annual Cocoa Beach Conference & Exposition, Cocoa Beach, Florida, January 12-16, 1997.4. Klosterman, D, R. Chartoff, B. Priore, N. Osborne, G. Graves, A. Lightman, G. Han, S. Pak, J. Weaver, "Structural Composites via Laminated Object Manufacturing543(LaM)," Solid Freeform Fabrication Proceedings, University of Texas at Austin, August, 1996.5. Kalmanovich, G., L. Dodin, S. Tu, "Curved-Layer Laminated Object Manufacturing," Proceedings ofthe 7th International Conference on Rapid Prototyping, University of Dayton and Stanford University, San Francisco, CA, April 1997.6. Kalmanovich,G., "Curved-Layer Laminated Object Manufacturing," Solid Freeform Fabrication Proceedings, University of Texas at Austin, August, 1996.7.

Kalmanovich,G., "Curved-Layer Laminated Object Manufacturing," Solid Freeform Fabrication Proceedings, University of Texas at Austin, August, 1996.7. Effects ofEnvironment on the Mechanical Behavior ofHigh-Performance Ceramics, A Compendium ofData Preparedfor Oak Ridge National Laboratory, University of Dayton Research Institute, Dayton, OR, July 1997.AcknowledgmentsThe work reported here was performed under a grant from the United States Defense Advanced Projects Agency (DARPA) and the Office of Naval Research (ONR). Dr. William Coblenz and Dr. Steven Fishman are the program officers.IllustrationsLoser" y.ypOO:::-----J---Optics//Layer outline and crosshatch~""'..----~...-... "....HOt Laminating rollerPart blockFigure 1 : Schematic of the LaM process (courtesy Relisys, Inc.

"....HOt Laminating rollerPart blockFigure 1 : Schematic of the LaM process (courtesy Relisys, Inc. [1D·544laser gantryLOM 1015solenoid switch ~airbrush nozzlesolvent vent ........1 - - - - ....- - - - -• • • • •l-----, reservoirplatformceramic tapeFigure 2 : Semi-automatic system for applying solvent to monolithic ceramic tapes.A.overlap zone: cut with fine crosshatchingpart"cube"B.overlap zone : burnished with defocused laser beamFigure 3 : Two strategies aimed at improving decubing by affecting the overlap zone between the waste cubes and the actual part: a) adaptive crosshatching and b) adaptive surface burnishing.

This diagram is not drawn to scale; with the materials used in this study, the layer thickness was the same as the cut width (0.25 mm).545Figure 4 : Monolithic SiC parts made on a LOM2030 using the adaptive crosshatching software.Figure 5 : Close-up of surface of body armor piece made with adaptive crosshatching.546Figure 6 : Cross section of reaction bonded, monolithic SiC bar made with the ceramic LOM process.Laser Cutting Monolithic Tape (requires low power to cut)\Strong InterfaceDecubingI...I Weak Interface//Fibers (require high power to cut)Figure 7: Illustration of origin ofLOM processing difficulties for separate layers of fibers and monolithic ceramic tapes.547wind SiC fiber tows on drumbrush with Furfural & methanol solutiondry, remove, flattenI .Ivacuum-bag, B-stage in oven at 105°C, 30 minutesstack fiber layer onto SiC tapeFigure 8: SiC tape-preg fabrication process.Figure 9 : 1/5 scale, turbine engine seal built from 7 layers of SiC tape-preg.

As it appears, this part is fully cured but has not been pyrolyzed or reaction bonded yet.548Figure 10 : SiC fiber/furfural prepreg cut with a copper vapor laser. The fiber diameter is approximately 15 flm and the cut channel is approximately 80 flm in width. High quality laser cutting is characterized by the vertical walls, clean fiber ends, and minimal burn back zone.0° fiber layerceramic tape ---..layer90° fiber layer250 flmFigure 11 : Photomicrograph of green SiC/SiC composite made with "TTLOM" test bed. The sample has been fully cured, but not fully pryrolyzed.549550.

~..... ''''...,cyde 2 (lOwe. 5 min,3~'rapidhea:lIp";;l:le(sanpleplac<din"'·"'I'~d..tI05O~C)Factors Involved in Tape CastingSlurry composition and viscosityDoctor blade settingCasting speedDrying conditionsConclusionsHNgiass formulation can be cast into useable tapes for lamination on the LOMHAlgIass tapes are appropriate to form LOM shapes with acceptable adhesion between la:yersA 3;.0 bone shape was successfully produced from a CAD file taken from a medical CT ScanHigh temperature post-processing techniques were successfully used to densify the LOM bones without distorting the complex shape.Formulation Selection3: I ratio was selected based on compressive strengthHowever, FTIR analysis of sintered specimens revealed the presence of a reaction by-product, Tricalcium phosphate (TCP), at a level of IO%. Becanse TCP at this level is resorbed in the human body, the material system cannot be considered nonresorbable in a strict medical sense.

Becanse TCP at this level is resorbed in the human body, the material system cannot be considered nonresorbable in a strict medical sense. Despite the presence ofTCP, the Current 3:1 HAfglass system was adopted to serve as a model system for use on the LOMStep 3: Tape CastingSoh'ent lBinder !~:':>:~:~f- i-i-. ; : .HA and GtllS.'iPowder MixingMiUing'with Sohl:nt, 24 hrs,Millingwitb Binder,24 Ms.Tape CastingConclusions (cont.)The mechanical properties of the specimens made with the best HA/glass fOImulation(3:1 ratio) were lower than that of cortical bone and pure sintered HA by a factor oftwo or three.The mechanical properties are expected to rise with optimization ofthe yanous processes involved, particularly post-pressing (elimination of interlammar yoids) and sintering.Ovem1l feasibility of producing bioceramic bone shapes from LOM has been established..

tomated Fabrication of Ceramic Components from Tape-Cast CeramicBrian B. Mathewson l Arthur H. Heuer3, Wyatt S. Newman2 , James D. Cawley4,ABSTRACTThis paper describes a machine and process for automated fabrication of functional 3-D laminated engineering components, ceramics in the present example. A laser cuts successive layers of a part derived from a CAD model description out of unfired tape-cast ceramic sheets vacuum-clamped to an x-y sled. A material-handling robot uses a selective-area gripper to extract only the desired part outlines from the surrounding waste material, then stacks the slices to build the part. This system design enables rapid manufacture of functional engineering components with arbitrarily complex internal and external geometries from virtually any material available in sheet form.I. IntroductionThis paper describes the results of our investigation of a machine and process for automated fabrication of laminated engineering components.

IntroductionThis paper describes the results of our investigation of a machine and process for automated fabrication of laminated engineering components. The purpose of our system is to enable custom manufacturing of functional 3-D components with arbitrarily complex internal and external geometries.We examined four key areas: testing a novel means of manipulating individual laminae of arbitrary shapes; determining the achievable precision of robotically-assembled laminated components; determining the dimensional tolerance achievable by precompensating for part shrinkage that will occur during post-processing; and testing the mechanical strength of ceramic parts so produced. The first two are described here; structural performance and materials processing issues are presented in a separate publication within these proceedings [1].System OverviewThe CAM-LEM approach is illustrated in Fig. 1.

1. A part originates from a computer description, which is analyzed to decompose the part into boundary contours of thin slices. These individual slices are laser cut from "green" sheet stock per the computed contours. The resulting part-slice regions are extracted from the sheet stock and stacked to assemble a physical 3-D realization of the original CAD description. The assembly operation includes a "tacking," procedure that fixes the position of each sheet relative to the pre-existing stack.

The assembly operation includes a "tacking," procedure that fixes the position of each sheet relative to the pre-existing stack. After assembly, the layers arelCAM-LEM Inc., Cleveland, Ohio.2Associate Professor, Department of Electrical Engineering and Applied Physics, Case Western ReserveUniversity, Cleveland, Ohio.3Kyocera Professor of Ceramics, Department of Materials Science and Engineering, Case Western ReserveUniversity, Cleveland, Ohio.4Great Lakes Associate Professor of Ceramic Processing, Department of Materials Science andEngineering, Case Western Reserve University, Cleveland, Ohio.253"laminated" by warm isostatic pressing (or other suitable method) to achieve intimate interlayer contact, promoting high-integrity bonding in the subsequent sintering operation. The laminated "green" object is then fired (with an optimized heating schedule) to densify the object and fuse the layers (and particles within the layers) into a monolithic structure.

The laminated "green" object is then fired (with an optimized heating schedule) to densify the object and fuse the layers (and particles within the layers) into a monolithic structure. The result is a 3-D part which exhibits not only correct geometric form, but functional structural behavior as well.As in all Rapid-Prototyping or Solid-Freeform-Fabrication systems, temporary support structures are required when forming surface contributions with downward-pointing surface normals (e.g., cantilevers, "ceilings," "stalactites," etc.). The CAM-LEM process uses "fugitive" materials for support of green assemblies. Fugitive materials are laser cut from sheet stock, generating shapes complementary to respective desired part cross sections. Fugitive materials and engineering materials are cut alternately, enabling construction of layered assemblies of solid blocks with spatially-varying material properties.

Fugitive materials and engineering materials are cut alternately, enabling construction of layered assemblies of solid blocks with spatially-varying material properties. During postprocessing (firing), the fugitive material bums out and the engineering material fuses into a structural solid.The results described here were obtained using an experimental system developed at Case Western Reserve University. Components ofthe system include: a Coherent Model 42 50-Watt CO2 laser; a NEAT 100mm by 100mm travel x-y sled; a four-axis Seiko XY-2000 robot; a Sun SparcStation IPC host computer; a VME-based multiprocessor system for real-time control; a variety of pneumatic actuators and controls; and an aluminum-framed Plexiglas safety enclosure. We integrated these elements into a system suitable for performing our feasibility evaluation experiments.Figure 2 is a photograph of a variety of materials cut with the experimental system.

We integrated these elements into a system suitable for performing our feasibility evaluation experiments.Figure 2 is a photograph of a variety of materials cut with the experimental system. The part outline shown is a standard mechanical-test part, which we used for strength testing of ceramic assemblies. This photograph also illustrates the extensibility of CAM-LEM to a wide variety of material choices. The materials shown are, from left to right: paper, tape-cast alumina, cardboard, iron-nickel powder in a BASF binder, Plexiglas, and Styrofoam. While these materials will utilize different bonding techniques, the CAM-LEM software and machine design apply universally.Figure 3 illustrates the precision and complexity manageable with the CAM-LEM approach. The object shown is comprised of 307 layers of paper. The source CAD file (generated by Laser Design, Inc., using a 3-D laser scan) originated in "STL" form--the de-facto standard for rapid prototyping [2]. CAM-LEM, Inc.

The source CAD file (generated by Laser Design, Inc., using a 3-D laser scan) originated in "STL" form--the de-facto standard for rapid prototyping [2]. CAM-LEM, Inc. 's software successfully translated the file into slices and contours, and the experimental CAM-LEM system generated each slice; assembly was performed manually. It is the objective of CAM-LEM, Inc. to create a commercial system capable of fabricating objects with the precision and complexity exemplified by Fig. 3. Such a system should be fully automatic, and capable of constructing objects from a variety of engineering materials. Assemblies so constructed will be transformable into fused solids, where the integrity of inter-laminar bonding rivals that of the bulk material properties themselves.II. ResultsCAM-LEM, Inc. in cooperation with CWRU investigated several key technological issues enabling the realization of such a system.

ResultsCAM-LEM, Inc. in cooperation with CWRU investigated several key technological issues enabling the realization of such a system. The machine-design issues of achieving the demanding material handling and the precision automatic layered assembly, are described here.254Computer ModelContour RepresentationSlice CuttingStackingLaminationFiringFinished ComponentFigure 1: The CAM-LEM Process255I256Material HandlingUnder the CAM-LEM technique, applied to ceramics, cross-sectional shapes of individual laminae are cut from sheet material, robotically assembled and tacked into 3-D objects, laminated into structurally-sound green (unfired) solids, and sintered to densify the ceramic particles into a monolithic body. Cutting desired regions from individual sheets offers prospective advantages over techniques which cut, deposit or fuse material directly on top of a subassembly.

Cutting desired regions from individual sheets offers prospective advantages over techniques which cut, deposit or fuse material directly on top of a subassembly. These techniques include stereolithography [3], selective laser sintering [4], three dimensional printing [5], and laminated object manufacturing [6]. By cutting laminae individually, the geometric formation process is fully decoupled from the material processing steps, in sharp contrast to most existing SFF techniques, thus obviating compromises in either operation. Conceptually, the CAM-LEM approach is most similar to the Helysis system of "Laminated Object Manufacturing" (LaM) [6], in which laminated objects are formed from stacks of laser-cut sheet material. The primary difference between LaM and CAM-LEM is that the LaM system stacks sheets first, then cuts outlines, whereas CAM-LEM cuts first then stacks.

The primary difference between LaM and CAM-LEM is that the LaM system stacks sheets first, then cuts outlines, whereas CAM-LEM cuts first then stacks. Due to this difference, the CAM-LEM system more readily accommodates internal voids and multiple materials, but at the expense of more demanding material handling. In particular, the CAM-LEM approach introduces the additional complexity of extracting desired cross-sectional regions after cutting them from continuous sheets. To solve this problem, we have developed a novel selective-area gripper, capable of extracting geometrically complex regions of material from the cutting plane.We have successfully tested cutting and manipulating various materials including paper, cardboard, Coors Ceramics tape-cast ceramic, CWRU tape-cast ceramic, plastics, and tape of powdered metals. Our tests have shown that the gripping operations are virtually independent of the type of material used.

Our tests have shown that the gripping operations are virtually independent of the type of material used. We used identical functions, procedures, and parameters in picking up and stacking both the paper and ceramic tape, for example.The design of the cutting surface is important for obtaining high-quality cuts. Initially, a smooth aluminum cutting surface was used. However, parts cut on this surface exhibited poor edge quality (possibly due to reflections of the laser from the aluminum surface). Further, polymeric binder material vaporized by the laser would recondense on the cutting surface, further roughening the cut edges, and often causing the desired part to adhere to the table and/or to the waste material. In addition, parts placed loosely on the cutting surface were displaced by the laser's air-jet flow.

In addition, parts placed loosely on the cutting surface were displaced by the laser's air-jet flow. (The air-jet flow is necessary to protect the lens of the laser, and to help eject debris from the laser cuts, thus improving edge finish.)To prevent part displacement during cutting, the cutting surface plate was punctured with an array of holes, and a vacuum suction was applied to clamp the sheet material to the cutting surface. This approach prevented part slippage but exacerbated the recondensation and adhesion problem. Alternative surfaces of wire mesh, wire cloth, acrylic plates, ceramic mesh, and others were tested. In these cases, rough surfaces provided too little vacuum clamping force and smooth surfaces had problems with material condensation and adhesion.The most successful cutting surface tested thus far consists of an aluminum "honeycomb" material commonly used in aerospace applications requiring strong, lightweight structures.

This material provided good vacuum clamping and little material condensation and adhesion. The honeycomb cutting table structure provides a sufficiently high density of support for small part It also provides adequate ventilation beneath the laser cutting surface for debris feature sizes. released through laser cutting since the cell wall thickness is smaller than the laser beam diameter. Prior tests showed that the laser does not cut aluminum at the power levels required to cut slice materials, and that few materials tested thus far adhere to aluminum. This combination of factors257results in a strong, robust, and reliable solution to the cutting table design. A ceramic grid cutting surface with which we had some success can be seen in Figs. 4 and 6.Figure 4 is a photograph documenting the success of our selective-area gripper.

A ceramic grid cutting surface with which we had some success can be seen in Figs. 4 and 6.Figure 4 is a photograph documenting the success of our selective-area gripper. On the left side of this photo, a top view of the robot's gripper is visible, as well as the outline of a complex part cut from alumina tape, extracted, and stacked by the robot. The part shown is actually a 2-layer assembly, stacked with high precision. On the right side of the photograph is the x-y sled's cutting surface. Alumina tape stock was placed on this surface, laser cut, and the desired part area was extracted by the gripper. The complement of the desired part shape can be seen clearly, including the small circular cutouts, which were left behind by the gripper when the desired area was extracted.Figure 4: Selective-area extraction of complex shape.Assembly Precision EvaluationA second key feasibility evaluation was to determine the achievable precision of robotically assembled laminated components.

To quantify possible inaccuracies due to part slippage during gripper manipulation, we devised an automated test.The computer commanded the robot to repeatedly pick up a ceramic specimen (a z-motion), translate the part approximately the distance required to perform assembly (a y-motion), then return the specimen to its original position (reverse y and z motions) and release the part on the table. This sequence of operations simulates what happens in an actual cutting and stacking operation. The dog bone shape employed for this test, later used for strength-test experiments, offered sufficient geometric complexity (curved edges, both wide and narrow sections, and internal cutouts) to comprise a realistically challenging evaluation. Part position and orientation were computed by calculating the part centroid and orientation from images acquired by a CCD camera mounted above258IJVCU.LJ.'-HJ. of the part.

Part position and orientation were computed by calculating the part centroid and orientation from images acquired by a CCD camera mounted above258IJVCU.LJ.'-HJ. of the part. After each manipulation cycle, the position of the part and its anglecamera were recorded to determine placement repeatability.Placement Error, Y-Axiso510152025Iterations Placement Error, Rotation0 .1 . - - - - - . - - - - - . - - - - - - - - - . - - - - - . - - - - - - - .0.08 (3) c3 0.06 eO.04w 0.0200510152025IterationsPart placement error during selective-area gripper handlingplots show the position of the part centroid location and part orientation between each iteration. The microns and 0.021 degrees, 0.021 degree rotation about the center of the part results in a displacement of the5 shows the results of 25 iterations of the test of part placement.standard deviation of the readings for y and the angle were respectively.

outerof the test specimen by 14 microns.these results we conclude that our selective area gripper is capable of moving part fact, handling errors are within the resolution limit of the Seiko measurements only establish a conservative upper bound on the achievablehigh repeatability.robot, soFurther tests were performed on a robotically-stacked rectangular assembly, shown in Fig. 6. cross section, 24mm on a side, was laser cut froin Coors alumina tape, lifted by the layers of 61 Omm-thick roughness of robotically-stacked assemblies has not yet been formally our experimentsandonto the assembly.

The assembly shown includesWhile the~""'J.LJ.JlJ.""""", we subjectively conclude that the stacking accuracy exhibitedexcellent assembly precision.have demonstrated the ability of the CAM-LEM system to create complex parts from paper models exhibit precision cOinparable to other SFF We have also shown how the CAM-LEM process can utilize other sheet materials, such bulkindustry-standardfiles.ascreatecomponents with structural properties rivalingmaterial properties themselves. The cut-then-stack approach facilitates the formation of interior voids and channels without the need to remove waste manually. Further, fugitive material added to the assembly can support cantilevered structures. Further work will investigate the ability of a 5-axis cutting system to cut edge tangents to optimize build time and surface finish by using thicker material layers.

Further work will investigate the ability of a 5-axis cutting system to cut edge tangents to optimize build time and surface finish by using thicker material layers. We conclude that a CAM-LEM Solid Freeform Fabrication system is feasible and are pursuing development of a fully-functional prototype.Figure 6: Cut and Robotically Stacked SquaresAcknowledgmentsThe authors acknowledge the significant contributions of Fong Choon Chiak, Sangeun Choi, Yong Zheng, Paul Wei, and Dr. Z. E. Liu towards the CAM-LEM project. This work was supported and funded by: the U. S. Army Research Office, contract #DAAH04-94-C-0055; National Science Foundation, grant #DMI-94-20373; National Science Foundation, Young Investigator Award, grant #IRI-92-57269; Office of Naval Research, contract #N00014-95-1-0107; the Cleveland Advanced Manufacturing Program (CAMP); and the CWRU Center for Automation and Intelligent Systems Research (CAISR).References [1]1. Cawley, P. Wei, Z. E. Liu, W. S. Newman, B. B.

Cawley, P. Wei, Z. E. Liu, W. S. Newman, B. B. Mathewson, A. H. Heuer, "AI20} Made by CAM-LEM Technology", Solid Freeform Fabrication Symposium Proceedings, Austin, Texas, August 1995.[2]J. M. Pacheco, "Rapid Prototyping," Technical Report MTIAC SOAR-92-01, Manufacturing Technology Infonnation Analysis Center, 10 W. 35h St., Chicago, IL 60616, Cleveland, Ohio, September 1992.[3]P. F. Jacobs, Rapid Protoyping and Manufacturing, Society of Manufacturing Engineers, 1992.[4]H. L. Marcus, J. J. Beaman, J. W. Barlow, and D. L. Bourell. "From Computer to Component in 15 Minutes: The Integrated Manufacture of Three Dimensional Objects," JOM, 42(2):8, 1990.[5]E. Sachs, M. Sima, and P. Williams. "Three Dimensional Printing: Rapid Tooling and Prototypes Directly From a CAD Model," J. Eng. Ind., 114(4);481, 1992.[6] M. Feygin and B.

Sachs, M. Sima, and P. Williams. "Three Dimensional Printing: Rapid Tooling and Prototypes Directly From a CAD Model," J. Eng. Ind., 114(4);481, 1992.[6] M. Feygin and B. Hsieh, "Laminated Object Manufacturing (LaM): A Simpler Process," in Solid Freeform FabricationSymposium Proceedings, Austin, Texas, August 1991, p. 123.260.

This paper uses a goose neck door hinge component to illustrate the effects of lattice structure optimization when stiffness criteria drive part design. The effect of intermediate lattice cell density parameters on resulting lattice configurations from automated lattice structure optimization are studied and it is found that the compliance of the model depends upon the range of intermediate density elements present. The paper then compares the effect of a displacement constraint on optimized weight from rib-stiffened and lattice-stiffened shell models. It is shown that optimized weight results from the lattice configuration depend on part stiffness requirements. The results show that lattice structures can be successfully implemented in weight-critical components where relaxation in the displacement constraint is acceptable.I.

The results show that lattice structures can be successfully implemented in weight-critical components where relaxation in the displacement constraint is acceptable.I. INTRODUCTIONAdditive manufacturing (AM), also popularly known as 3D Printing is a layer based manufacturing approach in which a complete three-dimensional part is fabricated by adding materials layer by layer [1]. Due to this layer-based additive approach, parts with higher geometric complexity relative to conventional methods can potentially be fabricated. This capability provides the designer with higher design freedom to optimize the part design based on problem physics for optimum performance rather than being limited by manufacturing constraints [2].One of the capabilities enabled by additive manufacturing is hierarchical complexity, i.e., the ability to build features at multiple size scales: micro, meso and macro [1]. This paper focuses on meso-scaled repetitive cellular structures like lattice structures.

This paper focuses on meso-scaled repetitive cellular structures like lattice structures. These cellular structures can be designed in such a way that they can be used to fill certain regions of a geometry. From a structural perspective, the main advantage provided by a lattice structure is high strength to weight ratio when being used to target mass reductions [3]. Also, lattice structures provide good energy absorption, thermal and acoustic characteristics [1]. They can also be used as support structures in several additive processes with minimum support material usage [4].Lattice structures are inspired from the unique repetitive arrangement of atoms and the bonds between them in crystalline solids. Similarly, in this case, these structures are modelled as interconnected struts between nodes in terms of unit cells and these unit cells are repeated in three- dimensional space to create a truss like structure [5].

The main parameters that are to be controlled in a unit cell are the end diameters of each lattice member and the cell type. Though, there are many types of unit cells, the tetrahedral mesh used in the hinge model in this work© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2123 1generates lattice beam elements on the edges of relevant tetrahedral intermediate density elements to form unit cells..This work uses a goose neck door hinge component to illustrate the effects of lattice structure optimization when stiffness criteria drive part design. The lattice optimization is executed using an approach driven by a topology optimization algorithm. During topology optimization, the density of elements are discretized as either 0 (void) or 1 (solid). Whereas, when a lattice optimization is performed, the intermediate density elements are converted into lattice structures.

Whereas, when a lattice optimization is performed, the intermediate density elements are converted into lattice structures. Due to the porous nature of these cellular structures, compliance of the model tends to increase upon their inclusions. The first half of the paper illustrates the effect of intermediate cell density settings on compliance of the model and the second half of the paper compares the effect of a displacement constraint on the mass of rib stiffened and lattice stiffened models.II. BACKGROUNDPrevious work performed in the fields of lattice structures, topology optimization, latticeoptimization, and the effects of lattice optimization on part stiffness is discussed in this section.A. LATTICE STRUCTURESDue to advancements in AM, there has been a considerable increase in research in the area of cellular structures and especially lattice structures. A significant amount of work has been done to investigate the applications and properties of lattice cells.

A significant amount of work has been done to investigate the applications and properties of lattice cells. Stucker, et al. [6], has shown that periodic lattice filters can be used efficiently in casting of molten metals as a replacement to ceramic foam filters. Additionally, due to its excellent thermal properties, heat exchangers are modeled after a tetrahedron diamond lattice design by Heidrich et al. [7] and in the field of metal AM, lattice cells are widely used as support structures leading to minimal support material usage (Hussein et al.) [8].Extensive research has also been done on lattice cell parameters and properties. Most commonly, strut thickness is considered as a key design variable in a lattice unit cell. Whereas, the work by Tang et al. [5] shows that the orientation of lattice cells plays an important role on structural properties. Iyibilgin et al.

Whereas, the work by Tang et al. [5] shows that the orientation of lattice cells plays an important role on structural properties. Iyibilgin et al. [9] observed that lattice structures built using fused deposition modeling process has higher strength compared to specimens with same porosity built using the sparse and sparse-double dense styles. Another work by Maskery et al. [10] demonstrates through experimental testing that the mechanical properties of latticed parts are a function of unit cell size. In cellular structures, the octet truss lattice configuration is a widely studied cell structure. Deshpande et al. [11] have analyzed the mechanical properties of an octet unit cell by considering it as struts that are pin jointed at its vertices and have found good agreement in results between analytical predictions and FE calculations.B.

TOPOLOGY OPTIMIZATIONA topology optimization problem is formulated such that optimal spatial distribution of material is obtained for a given set of loads and boundary conditions to minimize or maximize a certain objective function [12]. Topology optimization is performed to obtain a concept design that usually requires further fine tuning by shape and sizing optimizations. Several optimization algorithms are developed for this purpose. Initially in 1988, a homogenization method based on square and rectangular holes was introduced by Bendsoe and Kikuchi [13] and it was further developed into the SIMP method (Solid Isotropic Microstructure with Penalization) by Rozvany, who has performed a detailed evaluation of available topology optimization algorithms in his work [14].© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2124 2C. LATTICE OPTIMIZATIONLattice optimization has been implemented as an extension of topology optimization.

LATTICE OPTIMIZATIONLattice optimization has been implemented as an extension of topology optimization. The main task involved in lattice optimization is to develop an algorithm for designing lattice structures. Several methods have been proposed for designing these cellular structures. For instance, Nguyen et al. [15] proposed a two-step approach for designing conformal lattice structures (CLS) based on a heuristic which assumes, stress distributions are similar in CLS and in a solid body of same shape. As an alternative to the time consuming traditional CAD software, McMillan et al. [4] have developed a programmatic method for generating lattice structures directly in the STL format and in order to deal with computational and storage complexities in CAD systems caused by large conformal truss structures, a hybrid modeling method was developed by Wang et al. [16].

[16]. Further, a size matching and scaling method was developed by Chang et al.[17] for designing meso-scaled lattice structures. Hao et al. [18] have established an image based algorithm for developing periodic lattice structures and have investigated its manufacturability using selective laser melting (SLM) process. A comparison was done between uniform voxel based approach and conformal lattice approach for constructing lattice structures by Park et al.[19]. Finally, another comparative study was done between Particle Swarm Optimization (PSO) and least-squares minimization (LSM) for designing cellular structures by Chu et al. [20]. Later, it was found that PSO was more effective in searching large design spaces whereas LSM converged more quickly than PSO.This work uses the Altair OptiStruct Optimization program [21] for lattice optimization. Specifically, the work uses OptiStruct version 13, which is the initial lattice optimization release and has certain capability limitations.

Specifically, the work uses OptiStruct version 13, which is the initial lattice optimization release and has certain capability limitations. Some of the current limitations of the program include: lattice strut mesh dependency, lattice structures limited to circular cross section, and an inability to remove lattice members below a threshold diameter limit.D. EFFECT OF LATTICE OPTIMIZATION ON PART STIFFNESSThis work illustrates the dependence of structural compliance on the extent of lattice structures present in a part. Similarly, Rosen et al. [22] has demonstrated that stiffness of octet- truss structure decreases with an increase in the number of cells and Cerardi et al. [23] observed the inverse relationship between mechanical properties (stiffness and tensile strength) and the porosity rate (lattice cells) in three different structures. These results are in agreement with an investigation performed by Sudarmadji et al.

These results are in agreement with an investigation performed by Sudarmadji et al. [24] who demonstrated the inverse relationship between stiffness and porosity in functionally graded scaffolds using 13 different polyhedral configurations.III. LATTICE STRUCTURE DESIGN OPTIMIZATIONThis section provides the model details and its topology and lattice optimization results.Further, the effects of intermediate density range and porosity parameter value are discussed.A. HINGE MODEL DETAILSIn this work, a goose neck door hinge is used for illustrating the effects of lattice optimization. This component is used to open doors in aircraft and automobiles where an abrupt throw out action is necessary to clear an outwardly curved surface or other obstructions [25].

This component is used to open doors in aircraft and automobiles where an abrupt throw out action is necessary to clear an outwardly curved surface or other obstructions [25]. Figure 1 shows a conventional design of the door hinge.© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2125 3Figure 1: Conventional design of goose neck door hingeThe key area to focus is the location of placement of stiffeners along the hinge neck. The component is assumed to be made of aluminum. The blue region is the design space and the green region called “Lug” is non design space as shown in Figure 2. The base of the hinge is constrained by means of six fasteners to the door. Door rotation occurs about a pin through the lug. The lug is subjected to four separate loading conditions as shown in the free body diagram in Figure 2. The horizontal load represents the door in an open position whereas the vertical load represents the door in a closed position.

The horizontal load represents the door in an open position whereas the vertical load represents the door in a closed position. The other two side loads are to account for inertial forces.OpenCloseSide1Figure 2: Design Space and Free body diagram of Goose neck door hingeB. HINGE TOPOLOGY OPTIMIZATIONThis work is an extension of work done by Taylor et al. [26]. Here, Altair’s HyperMesh software is used for preprocessing and OptiStruct solver is used for optimization. The objective of topology optimization is to obtain the stiffest material distribution for the applied loads and© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2126 4Side2boundary conditions as shown in Figure 2. The OptiStruct solver is based on DRCO (Design variable, Response, design Constraint and Objective function) approach.OptiStruct uses a density based approach to solve topological optimization problems [27].

Under this method, the element density is used as a design variable that should ideally take a value of 0 (void) or 1 (solid), which would make it a discrete variable leading to higher computational needs. In order to bypass this, element density is treated as a continuous variable between 0 and 1 and any intermediate value represents fictitious porous material, which is not meaningful in conventional materials. Hence, these intermediate density values are penalized, forcing the final design toward element densities of 0 and 1.The hinge optimization study uses volume fraction and weighted compliance responses. Volume fraction refers to the percent of initial design space to be maintained in the final solution and weighted compliance is the sum of compliance of four individual load cases. Both the responses are global and are defined for the whole structure.

Both the responses are global and are defined for the whole structure. A specific percent of volume fraction is used as a constraint and minimizing weighted compliance is the objective function used to obtain the stiffest configuration. Based on the optimization results for various values of volume fraction constraints, 40% is chosen for further analysis as it provides better feature definition compared to other models.Topology optimization provides an optimum design based on critical load paths which are often not manufacturable through conventional methods. Without manufacturability constraints, the optimized configuration, shown in Figure 3, has a hollow interior, which is effective in resisting both the bending and torsion conditions induced by the load cases.Figure 3: Optimization result for 40% volume fraction constraintimprove manufacturability with conventional machining operations [28] but this constraint penalizes the torsional stiffness of the optimized design.

The optimized result with the draw direction constraint is shown in Figure 4. This configuration resembles an I beam section, which resists bending during the open and close loading conditions better than other beam cross sections but is less effective at resisting side load-induced torsion than the previous hollow section configuration.A draw direction constraint can beimposed forlineartool accessto© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2127 5Figure 4: Optimization result – 40% volume fraction with draw direction in Z axis and no holes constraintC. HINGE LATTICE OPTIMIZATIONLattice optimization provides a method to create solid components combined with lattice leverage additive manufacturing capabilities.

HINGE LATTICE OPTIMIZATIONLattice optimization provides a method to create solid components combined with lattice leverage additive manufacturing capabilities. Latticestructures, which can potentially optimization is a two-step approach [29] as shown in Figure 5.Regular Topology OptimizationLattice OptimizationInclude intermediate density elementsConvert intermediate density elements into lattice struts •Size lattice membersFigure 5: Lattice Optimization Flowchart.Whereas in regular topology optimization, the intermediate density elements are treated as fictitious material and are penalized into voids and pure solids, in lattice optimization, these intermediate density elements are converted into lattice structures using a two phase process. During the first phase, regular topology optimization is carried out, except that intermediate density elements are not penalized and are retained within the model [29].

During the first phase, regular topology optimization is carried out, except that intermediate density elements are not penalized and are retained within the model [29]. The range of intermediate density elements to be present can be controlled and it will be discussed in the next sub-section. In the second phase, the omitted intermediate density element edges are converted into lattice rod elements and the end diameters are sized based on a stress constraint for further fine tuning. In order to perform lattice optimization, the LATTICE command is included in the DTPL bulk data entries as shown in Figure 6.Figure 6: Bulk data section in .fem file© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2128 6Inclusion of the LATTICE command allows the user to specify the cell structure, lower and upper bound for intermediate densities and stress constraint values for lattice sizing.

Density values below the lower bound “LB” will be converted into void and values above the upper bound “UB” will be converted into solids. Elements between LB and UB are converted into 1D simple beam elements (Type ROD) with diameters proportional to the density of the intermediate density elements that were replaced [29]. Figure 7 shows the lattice optimized result and the sized lattice end diameters. The limitations of the initial lattice optimization implementation in OptiStruct version 13.0 can be seen in the mesh dependent lattice members and the inability to control the orientation and lattice cell configurations. Despite these limitations, this lattice optimization capability enables the component stiffness study executed in this work.Figure 7: Lattice optimized hinge and sized lattice end diameters.1.

Despite these limitations, this lattice optimization capability enables the component stiffness study executed in this work.Figure 7: Lattice optimized hinge and sized lattice end diameters.1. EFFECT OF RANGE OF INTERMEDIATE DENSITY ELEMENTSThe use of lattice structures can reduce component weight but it should be noted that lattice structures display lower stiffness per volume compared to fully dense material which increases the compliance. Hence, the amount of increase in model compliance depends upon the density range of intermediate elements present which are later converted into lattice structures. A stress constraint of 20 ksi is constantly used throughout this work for sizing the lattice diameters in the second phase of optimization. The increase in compliance relative to the range of intermediate density elements is shown in Figure 8.

The increase in compliance relative to the range of intermediate density elements is shown in Figure 8. The open and close loading conditions drive the design in very similar trends and hence the plot of compliance vs intermediate density element range is explained only for open loading condition.

The open and close loading conditions drive the design in very similar trends and hence the plot of compliance vs intermediate density element range is explained only for open loading condition. It can be seen from the graph that the compliance steadily increases as the upper bound of the intermediate density range increases, which means higher the upper bound the larger the number of lattice structures and higher the compliance.© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2129 7Percentage increase in compliance for Open load case80%70%60%50%40%30%20%10%0%10-30%10-50%20-50%40-50%10-60%30-60%20-70%Figure 8: Compliance Vs Intermediate Density Elements Range for Open Loading ConditionThe trend which is observed in Open and Closed loading conditions cannot be applied to the side loads because side loads puts the hinge under torsion whereas Open and Closed loading conditions causes bending.

A box type structure as shown in Figure 3 would prove efficient for resisting torsion but as the draw direction constraint is imposed in lattice optimization, material along the sides is removed and replaced with lattice structures that increase the compliance for the side loads. The magnitude of the increase remains more or less the same throughout the entire range of intermediate densities studied as shown in Figure 9. During Open or Closed loading, the lattice elements resist shear and during Side loading, they resist twisting. Both sets of load cases drives the design differently and it can be observed in plots Figure 8 and Figure 9.100%Percentage increase in compliance for Side load cases80%60%40%20%0%10-30%10-50%20-50%40-50%10-60%30-60%20-70%Figure 9: Compliance Vs Intermediate Density Elements Range for Side Loading Condition© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2130 810-80%10-80%2.

EFFECT OF POROSITY PARAMETERFrom the previous section, it is shown that compliance of the model depends upon the amount of lattice structures present, which is controlled by specifying the range of intermediate density elements. Another way to control the amount of lattice structures present is by using the porosity parameter [29]. The porosity parameter controls the penalization of intermediate density elements in the first phase of optimization. It is very similar to penalization which occurs during regular topology optimization except that intermediate density elements are retained in the model. Three options are available for this design optimization parameter: High, Medium and Low.If the Porosity value is High, then there is no penalty applied which leads to a larger amount of intermediate density elements in the first phase, resulting in a high volume fraction [29] of lattice structures in the final design.

It is to be noted that the model will have very high compliance due to the large amount of lattice structures. For a Medium porosity value, the penalty applied is 1.25 which leads to reduced lattice structure zones compared to the previous option. Both High and Medium options are preferred in applications where the component porosity is desired. For instance, in the case of a biomedical implants, porosity helps in the growth of tissues over the implants [22]. When the porosity value is Low, a natural penalty of 1.8 is applied which generates a further reduction in intermediate density elements in the first phase and leads to a design with very low lattice zones and mostly fully dense material distribution. This option is preferred in obtaining a stiffest design (compliance minimization problems).

This option is preferred in obtaining a stiffest design (compliance minimization problems). The variation in porosity in shown in Figure 10 for lattice optimized model with 10 to 80% intermediate density element range.High PorosityMedium PorosityLow porosityFigure 10: Variation in Porosity parameter: High, Medium and LowThough the above model has a very high intermediate density range of 10 to 80%, it should be noticed that the amount of lattice structures reduces drastically as we move from High to Low and hence there is a decrease in compliance as shown in Figure 11.

Also, for the low porosity option, the lattice structures are concentrated in the area of stiffeners as predicted by the topology optimized result from Figure 4.© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2131 990%e c n a80%70%i lp m o C n60%50%Openie s a e r c nI40%30%20%CloseSide%10%0%HighMediumLowFigure 11: Increase in compliance vs Load cases for various porosity optionsIV. SIZING OPTIMIZATION AND LATTICE STRUCTURESShell model created from the topology optimization result and other model variants withlattice structures along with its sizing optimization results are discussed in this section.A. HINGE SHELL MODEL DETAILSSizing optimization is required after obtaining an optimized topology result in order to determine geometrical parameters such as plate thickness for a 2D shell element or diameter for a lattice member is determined.

In sizing optimization, the structural connectivity of the model determined during topology optimization remains fixed throughout optimization process [30]. Whereas minimum compliance design in topology optimization adds material along critical load paths and provides the stiffest material distribution in the model, it does not take strength or stability requirements into consideration. Hence, a shell model as shown in Figure 12 is developed from the topology-optimized result and is size-optimized by taking strength, stability and displacement constraints into account.Figure 12: Shell model of the Hinge with StiffenersSimilar to the topology optimization model, the shell model is also constrained by means of six fasteners at its base and has the same loading conditions as shown in Figure 2. Gap elements are used at the base of this model as an interface element between two faces of the structure.

Gap elements are used at the base of this model as an interface element between two faces of the structure. Gap elements are one-dimensional elements defined by two nodes and are capable of transferring© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2132 10only axial forces (tension and compression) [31]. Here, initially the gap between the base of the hinge and the door is assumed to be closed and there is a load transfer only when the gap elements are in compression (closed condition).B. HINGE MODEL VARIATION AND SIZING OPTIMIZATIONIn order to understand the effect of the displacement constraint on optimized lattice structure, two other models were created apart from the Shell and Stiffener model (Model-I) shown in Figure 12. Model-II and Model-III consist of lattice structures in place of stiffeners, as shown in Figure 13, with the lattice structures imported from the lattice optimization result.

Model-II and Model-III consist of lattice structures in place of stiffeners, as shown in Figure 13, with the lattice structures imported from the lattice optimization result. Optimized lattice structures were used from the 10-80% intermediate density range configuration and the low porosity option. The main difference between Model-II and Model-III is the diameter of lattice structures. In Model-II, the diameter of lattice structures are the end diameters from the lattice optimized result whereas in Model-III, the lattice structures are divided into 3 sections along the neck of the hinge: forward, mid and backward region and corresponding sizing variables are created.

During the optimization, each section is sized individually so that all the lattice members in a single section have the same diameter.Figure 13: Shell model of the Hinge with Lattice StructuresIn order to transmit the load from the upper flange to lower flange through lattice structures, the nodes along edges of lattice structure are selected and connected to the flanges by using rigid RBE2 elements, which rigidly transmit displacement from one node to another.Similar to topology optimization, OptiStruct uses the DRCO approach for sizing optimization as well. The plate thickness of 2D shell elements and diameters of 1D lattice members are the design variables. Stress, displacement and mass responses are created. Stress constraints of max/min principal stress within ±40 ksi and a displacement constraint of 0.03 in. are applied for all three model variations.

Stress constraints of max/min principal stress within ±40 ksi and a displacement constraint of 0.03 in. are applied for all three model variations. For Model-III, tension and compression stresses of 40 ksi and 20 ksi, respectively, are given as constraints for sizing the lattice members. The buckling eigenvalue is constrained to be greater than 1.0 to ensure flanges and stiffeners remain stable under the design loads. Finally, minimizing the mass is the objective function for this problem and sizing optimization is performed.The resulting optimized design satisfies all stress, displacement, and stability constraints. Figures 14 -16 show Von Mises stress plots for all three models.. From the lattice stress plots, a small number of lattice members reach peak stress along the edges where they are tied to the© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2133 11shell flanges using RBE2 elements.

Figure 17 shows the resulting optimized shell thicknesses for all three models.Figure 14: Model-I Von Mises stress plot with deformationfor Open, Closed and Side load casesFigure 15: Model-II Von Mises stress plot for topology sized lattice membersfor Open, Closed and Side load casesFigure 16: Model-III Von Mises stress plot for three zone sized lattice membersfor Open, Closed and Side load casesFigure 17: Thickness plot for Model-I, Model-II and Model-III© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2134 12C. MASS COMPARISON FOR VARIOUS DISPLACEMENT CONSTRAINTSAs mentioned previously, though the use of lattice structures reduces the mass of component significantly, it also increases its compliance. In order to benefit from the usage of lattice structures, it is necessary to understand the relationship between mass and magnitude of the displacement constraint of the component.

In order to benefit from the usage of lattice structures, it is necessary to understand the relationship between mass and magnitude of the displacement constraint of the component. For this purpose, all 3 models were size-optimized for increasing displacement constraint values, holding stress and buckling constraints as previously set, and their optimized mass values were recorded. Figure 18 shows the optimized mass result versus displacement constraint magnitude.DISPLACEMENT VS MASSS S A M8.000 7.000 6.000 5.000 4.000 3.000 2.000 1.000 0.000Model-IModel-IIModel-III0 . 0 3 0 . 0 6 0 . 0 9 0 . 10 . 511 . 5DISPLACEMENT CONSTRAINTFigure 18: Displacement vs Mass for three model variationsTable 1: Mass (lbs) and %Lattice benefit for various Displacement constraintsTable 1 provides information about mass for all three model variations for increasing displacement values.

Here, %lattice benefit denotes the percentage reduction of mass by comparing Model-I and Model-II. With the initial tightly constrained diplacement values, essentially no benefit is obtained from the lattice optimization—the initial negative values are within numerical noise and model idealization error. Although the mass of Model-I and Model- II are almost the same initially, mass reduction increases to around 14% as the displacement constraint is relaxed. From Figure 18, the relation between displacement and mass is similar in© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2135 13all three design variants as expected, i.e. the mass of the component reduces if the displacement constraint is relaxed. The displacement vs. mass curve for Model I and II flattens after 0.5 in. displacement constraint magnitude because the thickness of the shell elements reaches a lower bound value of 0.05 in. during sizing optimization.

mass curve for Model I and II flattens after 0.5 in. displacement constraint magnitude because the thickness of the shell elements reaches a lower bound value of 0.05 in. during sizing optimization. From Figure 18, it can be seen that Model-III is comparatively heavier than the other models. Due to large number of lattice members, it is difficult to manually hand pick each lattice and create individual design variables. For this purpose, they were segregated into three sections and sized section wise. As the lattices are sized section wise based on stress, one or two lattice members determine the size of whole section which reduces the design freedom and leads to heavier weight.V. CONCLUSION AND FUTURE WORKThis work studied the effects of lattice structure configurations on component stiffness in a two phase effort. In both phases, a goose neck door hinge with four load cases is used for demonstrating the effects of lattice optimization.

In both phases, a goose neck door hinge with four load cases is used for demonstrating the effects of lattice optimization. In the first phase, topology optimization of the model provides an I beam configuration which resists bending during Open and Closed loading conditions and determines the location of stiffeners. Later, topology optimization is extended to lattice optimization and it should be noted that lattice structures are formed primarily in the place of stiffeners. Compliance of the model depends on the amount of lattice members present which can be controlled using two parameters: range of intermediate density elements and the porosity parameter. Figure 8 provides the relationship between the range of intermediate density elements and model compliance. It should be observed that model compliance increases as the density range increases.

Figure 8 provides the relationship between the range of intermediate density elements and model compliance. It should be observed that model compliance increases as the density range increases. Similarly, the porosity parameter can be used to control the amount of lattice structures present, which in turn determines the compliance of the model. There are three porosity options: High, Medium and Low. From Figure 8 and Figure 11, we can come to a conclusion that compliance of the model depends upon the lattice volume fraction, i.e., the higher the number of lattice members, the higher the compliance.The second phase of the work involves fine tuning of the design. To understand the relationship between lattice structures and displacement constraint, three shell model variations with stiffeners and lattice structures are created where the lattice members are imported from low porosity lattice optimization result.

These models are size optimized for increasing displacement constraint values and their corresponding mass values are recorded. From Figure 18 and Table 1, it can be noticed that mass of the component reduces as we increase the displacement constraint value and this trend is similar in all three model variations. About 14% mass reduction can be observed between the model with stiffeners and the model with lattice structures once we increase the displacement value. Based on the above observations, lattice structures can be implemented in solid models where low weight is preferred and at the same time relaxation in displacement constraint is acceptable.HyperMesh and OptiStruct version 13.0 was used for simulation purposes in this paper. The next version 14.0 has been released and work needs to be done to investigate the changes in results by using the enhancements of latest version.

The next version 14.0 has been released and work needs to be done to investigate the changes in results by using the enhancements of latest version. Some of the updated features in lattice optimization include use of tapered lattice members, lattice smoothing, and re-meshing features. Also, work is planned to study lattice structures designs in other applications where they can be used to reduce the component weight without affecting its stiffness. Future work will also examine methods to control the directionality, density, and cell configuration and to address the stress concentrations in the lattice structures which might lead to fatigue failure.© 2016 by Robert Taylor and Vignesh Dakshnamoorthy 2136 14VI. ACKNOWLEDGEMENTSThe authors wish to thank Altair Engineering, especially Blaise Cole, Hedison Mui, andChayan Basak, for their support of this work.VII. REFERENCES[1]I. Gibson, D. W. Rosen, and B. Stucker, Additive Manufacturing Technologies, Second Edi.

tomated Layer Decomposition for Additive/Subtractive Solid Freeform FabricationC. Chang, J. M. Pinilla, J.H. Kao, J. Dong, K. Ramaswami, F.PrinzDepartment of Mechanical Engineering, Stanford UniversityAbstractThe new additive/subtractiye technology (Shape. Deposition Manufacturing) enhances Solid Freeform Fabrication (SFF) capability inproducing near net-shape surface finish. This. technology also builds parts in fewer layers compared with conventional layered manufacturing technology. However, to decompose a part into freeform layers usually requires expensive geometric computation.. Also, to plan build. sequences often requires human intervention because ofthe complicated spatial relationships among the freeform layers. At present decomposition and build sequence planning are both performed by experienced designers/users.

At present decomposition and build sequence planning are both performed by experienced designers/users. In this paper, a novel decomposition approach based on surface splitting is proposed to facilitate .computation and planning of the additive/subtractive SFF processes. The results shown in this paper are from models with 3D planar geometry. Continuous effort is devoted into extending and implementing this new approach for models with 3Dfreeform.geometry.1. IntroductionTraditionalla.y~redmanufacturingtechniques have a shorter cycle time of realizing a part from a CAD model compared with that of typical machine shop practices.Cycletime is shortened mainly. because the technology employs a slicing strategy which greatly simplifies the planning task of buildingapart.•• lJnder this strategy, a given solidfi1odel is sliced horizontally into a set of planar layers, andthese planar layers are then built one at a time from bottom up.

In this case, planning the build sequence of a given model is no more than listing the sliced layers along the build direction, which is straightforward and can be automated easily.On the other hand, certain issues limit the application of this technology. First, parts built by this technology generally show stair steps on the layer-to-layer boundaries. This lack of C2 continuity on the part surfaces is undesirable especially for parts made of ceramic materials, because each stair step can serve as a crack initiation. site. In addition, the mechanical properties of parts will be affected by the bonding between layers. Clearly, reducing the number of layers by increasing the slicing interval can improve the mechanical properties. Nonetheless, the surface finish of the built parts may well become worse as the interval becomes larger.

Nonetheless, the surface finish of the built parts may well become worse as the interval becomes larger. As a consequence, a trade-off has to be made case by case.In contrast to traditional layered manufacturing technology, the new additive/subtractive SFF technology not only builds parts in fewer layers, but also achieves much finer surface finish. Instead of horizontal slicing, this new technology decomposes a solid model along a set of parting lines into freeform layers [3]. However, such decomposition usually requires expensive computation. In addition, the spatial relationship among these freeform layers is much more complicated than that among the planar .layers. As a result, the build sequence can not be determined without carefully examining the geometry of each freeform layer.

As a result, the build sequence can not be determined without carefully examining the geometry of each freeform layer. Therefore, the layer build sequence under this new technology is still completed mainly by experienced designers.111As shown later in this paper, a modified approach is developed for the additive/subtractive SFF technology in order to release designers from the time consuming task of decomposing a model and arranging build sequences. This new approach has been implemented successfully for models with 2.5D features. With limited testillg results. so far, it.is shown to be able to work for models with 3D freeform features as well.2. Surface SplittingCurrently decomposition of a given part is done in a "solid" level. That is, a givell •• solid model is split into smaller manufacturable solid portions.

Surface SplittingCurrently decomposition of a given part is done in a "solid" level. That is, a givell •• solid model is split into smaller manufacturable solid portions. An algorithm has been proposed [3] to into a set of freeform layers for Shape Deposition automatically decompose a 3D part In addition, all the feasible build sequences can be generated Manufacturing (SDM) processes. through a compact graph algorithm [6]. However, performing operationsin a "solid" level usually requires expensive geometric computation.To lower computationJoad, the. feasibility of utilizing surfaces of a solid to partition the solid into manufacturable freeform layers for SDM processes is investigated. Since the shape of any part is represented completely by its surfaces, theoretically we can realize a solid object according to its surfaces.

Since the shape of any part is represented completely by its surfaces, theoretically we can realize a solid object according to its surfaces. Simply stated, the object can be shaped out of a stock of material by performing machining operations on an surfaces of the object in such a sequence that each surface is machined without damaging any previously.built surface. With this concept, each surface on the solid serves as a build unit which "partitions" the whole solid. There is no need for splitting the solid.However, one critical manufacturing .characteristic has to be considered before the concept can be applied to SDM processes. In SDM processes, the build direction of a part is pre-selected and remains unchanged for all machining stages of the part, which means the cutters always access the target machining areas from the top along the build direction. As a result, some surfaces can not betnachined in a single step. As it can be seen in Fig.

As a result, some surfaces can not betnachined in a single step. As it can be seen in Fig. 1, we can not shape area B ofsurface S together with area A at one step. without damaging A.But surface S can be fl.llly shaped if we machine out area B first, deposit more part material on top ofB, and then machine out areaA. Under this reasoning, all surfaces of a part can be machined along the build direction as long as the surfaces .containing both undercut and non-undercut regions· are properly split into subsurfaces where each subsurface contains only either undercut or non-undercut region.s(a)(b)(c)(d)Fig.

S is a non-undercut surface ifN(u,v).b >= 0 for all (u,v) in S domain. S is a non-monotonic surface if it contains both undercut and non-undercut portions.NS5biMonotonic surface(a) Undercut surfaces: Sl, S3(b) Silhouette curves(c) Non-monotonic surface splittingNon-undercut surfaces: S2, S5 Non-monotonic surface: S4Fig.2 (a) Three surface types(b) Silhouette curves(c)Non-monotonic surface splittingSilhouette curveA silhouette curve is defined as the curve along which N(u,v).b=O. In other words, silhouette curves serve as the boundary between undercut and non-undercut portions of the surface. Thus one non-monotonic surface can be split iIltotwo or more monotonic (undercut or nOll-undercut) surfaces along.the silhouette curves as shown in Fig.

Thus one non-monotonic surface can be split iIltotwo or more monotonic (undercut or nOll-undercut) surfaces along.the silhouette curves as shown in Fig. 2(c).Overlapping surfacesFor any two surfaces, if there exist areas where a ray intersects both surfaces alohg the build surfaces.direction, these two surfacesSelf-overlapping surfaceIf a ray intersects a monotonic surface more than once along the builddirection,tllissurfaceis said to contain self-overlapping features. Fig. 3 illustrates the situation.2.2 Procedure descriptionFive steps are taken to split surfaces of a solid model: «& Generate silhouette curves on all the non-monotonic surfaces of the given solid model. Splitthese non-monotonic surfaces into monotonic surfaces through the silhouette • curves obtained in the previous step. Split self-overlap surfaces, if there exists any, into surfaces without self-overlap features [3].113Non-monotonic surfaceMonotonic surface..

Split self-overlap surfaces, if there exists any, into surfaces without self-overlap features [3].113Non-monotonic surfaceMonotonic surface.. Split overlapping surfaces so that each resulting surface is either totally overlapped or totallynot overlapped with other surfaces [3], then.. Mark each resulting surface as either an undercutor a non-undercut surface.A simple example is given in<Fig. 4.!(a)(b)Fig.3 Self-overlapping surfaceFig.4 (a) Generate silhouette curves (b) Six surface compactsAfter performing the five steps, the surfaces of the solid have been split into subsurfaceswhich serve as the basic build units.3. Surface Compact Graph ConstructionUp. to this stage,all ll1esurfaces of the target solid have been splitintosm.allersubsurfaces such that each subsurface can be processed in one additive/subtractive build cycle. The target solid can now be produced by shaping these.

The target solid can now be produced by shaping these. subsurfaces in certain buildsequelle~s.As mentioned earlier, a feasible build sequence is basically a spatial arrangement of subsurfaces which assuresth~ cutters will notrun into the already built subsurfaces while the target subsurfac~ is being machined. In order to identifysuch.seguences, the spatial relationships among allthese<.freeform subs~rfaces have to be examined and recorded.. A special graph strllctl1r~~an1edSompactGraph [4] is utilized to record these spatial relationships because it not only()ffers/i~conciseformatofkeeping these spatial information, but also provides useful. algorithms to/ explorealternativehllild sequences afterwards. Finally, a "preliminary surface. compact graph" is obtained after lhesubsurfaceentities and their spatial relationships are recorded through the Compact Graph structure.

Finally, a "preliminary surface. compact graph" is obtained after lhesubsurfaceentities and their spatial relationships are recorded through the Compact Graph structure. Again, the terminology and procedures are stated in the following sections.3.1 DefinitionsSurface compactA surface compact. represents a set of undercut or non-undercut surfaces which canibe machined/processed together in one SOM· cycle. In its simplest form, a surface compact is either an undercut surface or a non-undercut surface.S4Surface compact graphA graph in which each node represents one surface compact and each edge connecting two nodes indicates certain spatial relationship (either precedence or adjacency) between these two nodes.Precedence relationshipOne surface compact (or node) is "precedent" to another surface compact if there exists a ray which intersects both compacts along the build direction.

Thus if surface .compact A is "precedent" to surface compact B, compact A has to be built before compact B can be built. In other words,. the precedence relationship between two compacts represents the build order between them.Adjacency relationshipTwo surfacecornpacts (or nodes) are "adjacent" if they have common boundary.3.2 Procedure descriptionThree steps are taken to construct a preliminary surface compact graph: .. One node is created in correspondence with ..one.surface compact obtained through the surfacesplitting stage... Precedence relationship between each undercut surface compact and each non-undercut surface If there exists a. precedence relationship between the two surfacecompact is examined. compacts, a directed edge is created betweentheir associated nodes... The adjacency relationship between.any two surface compacts (or nodes) is then examined.

compacts, a directed edge is created betweentheir associated nodes... The adjacency relationship between.any two surface compacts (or nodes) is then examined. and an undirected edge is created if there exists adjacency relationship between the two compacts.the spatial relationship among all surface compacts is preserved in the resulting preliminary surface compact graph, as shown in the example below. (Fig. 5)After all surface compacts are processed through the three steps above,S6S6S6S5S4S5S4S5S4o nodePrecedence edgeAdjacency edge(a)(b)(c)Fig. 5 (a) Surface compacts (b) Surface compact nodes and prece<:iencee<:iges(c) Preliminarystirface compact graph1154. Surface Compact Graph ConsolidationThe preliminary surface compact graph obtained above represents the spatial relationships between every two surface compacts. However, redundant spatial relationships may well exist in this preliminary graph.

However, redundant spatial relationships may well exist in this preliminary graph. Because the computation cost of finding build sequences will be affected by these redundant spatial relationships, an intermediate step to consolidate the preliminary graph is taken before we search for the build sequences. At this stage, the preliminary surface compact graph is consolidated into a minimal surface compact graph.4.1 DefinitionsRedundant precedence edgeIf the precedence relationship represented by a precedence edge can be also represented through a combination of other precedence edges, this precedence edge is< called a "redundant" edge since it repeatedly expresses an existing precedence relationship.Minimal surface compact graphThrough the two steps described below, a preliminary surface compact graph can be consolidated into a more concise graph named minimal surface compact. graph.

graph. This minimal graph contains fewer nodes and edges but still describes the spatial relationship among the surfaces of a given solid model. Formally speaking, a minimal surface compact graph results when the nodes in a given preliminary graph can not be further merged and all the redundant precedence edges are removed. For instance, a solid with six surfaces can be represented by a minimal graph of four nodes and six edges rather than six nodes and thirteen edges as shown in Fig.

For instance, a solid with six surfaces can be represented by a minimal graph of four nodes and six edges rather than six nodes and thirteen edges as shown in Fig. 6.4.2 Procedure descriptionTwo steps are required to consolidate a preliminary surface compact graph into a minimal graph: e Any two adjacent nodes can be merged into a new node if no precedence edge exists betweenthem.e Locate and remove all redundant precedence edges in the graph with certain checking algorithm[5].After the adjacent node merge and redundant edge removal steps, a minimal surface compactgraph containing the least number of nodes and edges is formed.8684' = 85+8684'=85+86858483'=83+8483'=83+84Redundant edge82828181(a)(b)(c)Fig.6 (a)Preliminary graph (b)Node merge and redundant edge removal (c)Minimal graph1165. Build Sequence GenerationAs explained above, the minimal surface compact graph records all the build precedence relationships among nodes.

Build Sequence GenerationAs explained above, the minimal surface compact graph records all the build precedence relationships among nodes. Alternative build sequences can be explored by partitioning this minimal graph properly. Because the structure of a minimal surface compact graph is equivalent to that of a Compact Adjacency Graph (CAG) [4], the algorithms developed for CAG [6] can be also applied to a minimal surface compact graph. By employing these algorithms, the build sequence will be extracted automatically from the minimal graph as shown below. (Fig. 7)S4'=S5+S6S3'=S3+S4S4'=S5+S6S3'=S3+S4~'...........-., .../S2(a)(b)SlFig.7 (a) Minimal graph (b) Build sequence: Sl -> S2 -> S3' -> S4'6. ResultThe example part has twenty-four surfaces originally. Ten undercut surfaces and fourteen non-undercut surfaces are identified after performing surface splitting on the part.

ResultThe example part has twenty-four surfaces originally. Ten undercut surfaces and fourteen non-undercut surfaces are identified after performing surface splitting on the part. The preliminary surface compact graph contains twenty-four nodes representing these twenty-four surfaces. After graph consolidation, the resulted minimal surface compact graph contains only four nodes, which suggests this part can be built in four layers as illustrated in Fig. 10. In short, we first deposit support material and machine surface compact 81, then deposit part material and machine 82. More. support material is deposited on top of 82 and 83 area is machined. Finally, we deposit part material on top of 83 and machine 84 area.,~~----_._--(a)(b)Fig.9 (a) Example part (b) Decomposed surfaces117S3'S4'8483- .LY'?".~ ------ [~~~J838281------ \><=::\-, "~81i// ....... ...... i ,...... ....-~I: 1/(a)(b)(c)Fig.

...... i ,...... ....-~I: 1/(a)(b)(c)Fig. 10 (a) Minimal surface compact graph (b) Build sequence: 81->82->83->84 (c) Build sequence illustration7. ConclusionsInstead of decomposing a solid· model horizontally, the proposed decomposition strategy splits and distinguishes the. surfaces of the given solidinto undercut and non-undercut surface compacts. Based on the surface compacts, the compact graph algorithms can then construct build sequences of the given model automatically. Compared with traditional planar slicing apprpach, this new surface compact approach results in fewer build layers for 2.5D and 3D freeform solid models.8. AcknowledgementWe wish to acknowledge the financial support of DARPA under the 3D Reasoning projectand Prof. C. Sequin ofUC Berkeley for providing us the example part.9. References1. Harris L. Marcus and David L. Bourell. (1993) Solid Freeform Fabrication. Advanced materials and Processes, 144(3) pp28-352. Merz R., Prinz F.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAutomated layer identification in large area additive manufacturing (LAAM): A comparison of image thresholding and edge detection techniquesA. Wadidie*, G. M. Studer*, K. Villez**University of Maine Advanced Structures and Composites Center, Orono, ME, 04473**Oak Ridge National Laboratory, Oak Ridge, TN, 37830AbstractOur study aims to develop an automated method for identifying layers on images of 3D- printed walls from a LAAM printer, as manual identification is subjective and can be time- consuming. We applied three different image processing methods to identify edges between layers: simple thresholding, Otsu thresholding, and Canny edge detection. Otsu thresholding was found to be the most accurate and required minimal manual intervention.

Otsu thresholding was found to be the most accurate and required minimal manual intervention. From our study, we propose a new approach by going through essential steps for greater accuracy. This research demonstrates the feasibility of using computer-based methods to automatically identify layers in 3D printing, reducing manual time and effort and improving the strength and quality of 3D-printed parts.IntroductionThe advancements in Large Area Additive Manufacturing (LAAM) have ushered in a new era of efficient and cost-effective production of intricate three-dimensional structures. As LAAM structures are composed of layers of thermoplastic filaments, often planar, a significant factor in how a LAAM structure performs is the quality and consistency of adhesion at the layer interfaces. In most cases, it is very easy to identify the layer's boundaries on the surface of a 3D-printed object with human eyes.

In most cases, it is very easy to identify the layer's boundaries on the surface of a 3D-printed object with human eyes. However, the precise localization of all layers and boundaries on the surface of LAAM-printed objects is a complex and time-consuming task when performed manually, as a single print may contain hundreds of layers, each potential meters long.The development of an automated way to identify layers in 3D-printed walls is motivated by the need to quantify layer and boundary consistency while avoiding the time-consuming nature of manual identification. By accurately and automatically identifying layer boundaries, it becomes possible to detect anomalies as well as to analyze and optimize the printing conditions. In addition, the quality of layer adhesion and layer defects affect the overall quality, structural integrity, and many other mechanical properties of the printed object.

In addition, the quality of layer adhesion and layer defects affect the overall quality, structural integrity, and many other mechanical properties of the printed object. Manual identification of layers in 3D- printed LAAM walls is performed at the Advanced Structures and Composites Center (ASCC), but the process is subjective, tedious, error-prone, and not scalable for large-scale or time-sensitive projects. By developing an automated method, we aim to streamline the process and eliminate human subjectivity in identification. An automated method could be applied to a large number of1629images in a shorter time frame, making it more efficient and practical for industrial applications and real-time analysis and optimization.To develop an automatic layer identification method, our research is currently focused on state-of-the-art image processing techniques. We execute a comparative analysis of image thresholding and edge detection techniques.

We execute a comparative analysis of image thresholding and edge detection techniques. Three different methods are used to identify edges between layers in images of 3D-printed walls obtained from a LAAM printer. Initial analysis revealed that Otsu thresholding, a well-established image segmentation technique, exhibited the highest accuracy in identifying layer boundaries. Notably, it required minimal manual intervention, making it an attractive choice for practical implementation within LAAM processes.

Notably, it required minimal manual intervention, making it an attractive choice for practical implementation within LAAM processes. Each of these techniques is integrated into an image-to-layer data pipeline that is composed of a 4-step sequence of data processing steps.Two relevant developments that have influenced our research on layer identification in Large Area Additive Manufacturing (LAAM) are the Peregrine software [1]" and digital image correlation [2]." The Peregrine software is designed for in-situ quality control in powder bed systems and introduces AI-based software for defect detection in the top layer of a powder bed system, whereas digital image correlation addresses in-situ monitoring of residual thermal stresses and warpage in polymer composite printed components using digital image correlation.

We have taken inspiration from these studies and developed a new method for the specific purpose of automated identification of layer boundaries in LAAM structures, which has not been addressed before.Materials and MethodsData collectionWe used rectangular, vertically-printed 3D-printed walls as the primary material for our layer identification work (Figure 1). These walls were produced using a LAAM printer as part of an effort to produce material test specimens from ABS thermoplastic infused with carbon fiber, creating a black appearance, and were composed of 120 layers, each about 5mm high. The images were captured using a high-resolution camera at 8192 x 5464 pixels with overhead, uncontrolled overhead lighting from the facility ceiling. The wall itself was the primary feature in the camera scene and covered about 2/3rds of the image. We sampled from three different images with representative edge identification features.

The wall itself was the primary feature in the camera scene and covered about 2/3rds of the image. We sampled from three different images with representative edge identification features. The layers of the walls were generally printed imperfectly and included many variations and inconsistencies commonly encountered in LAAM. In addition, the various walls were printed at different speeds, resulting in further inconsistency.Other available LAAM materials included wood-fiber composites, which are morereflective but have a less smooth texture - these are considered for future work.1630Figure 1: 3D Printed WallData Preprocessing and TilingTo avoid image segmentation complications and focus on layer boundaries, the first step was to remove the background from our input images. As you can see in Figure 1, there are different types of objects around the wall, and the wall itself contains painted writing and ragged areas.

As you can see in Figure 1, there are different types of objects around the wall, and the wall itself contains painted writing and ragged areas. To tackle this, we initiated the process by cropping the images and selecting specific regions that could be accurately labeled, effectively removing the background and the irregular portions on the wall images (Figure 2).Each image is sliced into 128x128 tiles to be used when proceeding to the automated layeridentification for consistency in our metrics.Figure 2: Image cropping1631Manual edge identificationAfter tiling, the images are ready for manual and automatic labeling. Two different image manipulation and editing software tools were used - The first tool was Gimp, an open-source image editing tool [3], which we used to draw white lines along each edge between two layers, as shown in Figure 3.

As this was very time-consuming, we started using ImageJ [4], another open-source scientific image tool, which included more dedicated manual edge-detection tools. In ImageJ, we first had to convert the images to an 8-bit color format in order to apply a manual threshold to it. By adjusting the values depending on the intensity variations on the image, contiguous lines are possible. Figure 4 shows the same image manually labeled with ImageJ. After completing the manual labeling for each image, we would similarly tile the labeled masks into 128x128 tiles to match the image tiles.Figure 3: Manual edge labeling with GIMPFigure 4: Manual edge labeling with ImageJAutomated edge identification - OverviewThe pipeline in Figure 5 goes through the steps of our layer identification research.

Thesolid arrows are for the manual steps, while the dashed arrows are for the automated steps.1632Figure 5: Layer interface identification pipelineAutomated edge identification - Step 1: ThresholdingAfter manually labeling the sample images, we proceeded with the automatic labeling ofthe same images. We opted to employ three specific methods:Simple thresholding ● Otsu thresholding [5] ● Canny edge detection [6]First, we use simple image thresholding to identify the layers on the 3D-printed walls. Simple thresholding provided a straightforward and intuitive way to identify layers based on pixel intensities. The simple thresholding method converts the input image into a binary representation. A manually selected threshold value was applied to classify pixels as either edge or non-edge. Figure 6 shows the image sample, and Figure 7 is the resulting binary image.

A manually selected threshold value was applied to classify pixels as either edge or non-edge. Figure 6 shows the image sample, and Figure 7 is the resulting binary image. The edge pixels were colored in white, and the black areas were the non-edge pixels.1633Figure 6: Camera image sampleFigure 7: Method 1: Simple thresholding mask. White pixels are considered edge pixels. Black pixels are considered non-edge.For the second method, we utilized Otsu image thresholding, a well-known automatic technique for determining the optimal threshold value for each layer based on the histogram of pixel intensities (Figure 8). Otsu thresholding separates layer and edge pixels by finding the value that minimizes the variance between light and dark classes. It automatically analyzes the pixel intensity distribution, iterates through possible thresholds, and identifies the one that yields the clearest separation.

It automatically analyzes the pixel intensity distribution, iterates through possible thresholds, and identifies the one that yields the clearest separation. By incorporating Otsu thresholding, we achieved precise layer identification, enhancing our image analysis and enabling further investigation (Figure 8).Figure 8: Method 2: Otsu thresholding. Histogram of pixel intensities and typical results. Red pixels are considered edge pixels while blue pixels are considered non-edge pixels.1634Lastly, we employed Canny edge detection, a well-established edge detection algorithm developed by John Canny in 1986. Given its popularity and effectiveness in image processing and computer vision tasks, we selected Canny edge detection as our third method for layer identification. Canny edge detection identifies edges in an image based on the gradient of pixel intensities. It includes key steps for identifying edges in an image.

Canny edge detection identifies edges in an image based on the gradient of pixel intensities. It includes key steps for identifying edges in an image. It starts by applying a Gaussian filter to reduce noise and enhance edge quality. Then, the gradient of pixel intensities is calculated using the Sobel operator, providing the rate and direction of intensity change. Non-maximum suppression thins out edges by preserving only local maxima in the gradient direction. Finally, hysteresis thresholding connects potential edge pixels to form continuous strong edges. Canny detection is expected to produce accurate, well-defined edges while handling noise. By following these steps, we were able to extract the edges of interest for most of the images (Figure 9), providing valuable information for our analysis.Figure 9: Method 3: Canny edge detection.

By following these steps, we were able to extract the edges of interest for most of the images (Figure 9), providing valuable information for our analysis.Figure 9: Method 3: Canny edge detection. White pixels correspond to identified edgeswhile black pixels are considered as non-edge pixels.For each image, the three methods, simple thresholding, Otsu thresholding, and Canny edge detection, are applied in parallel. This approach allowed us to explore different techniques and assess their utility for layer identification purposes.Automated edge identification - Step 2: ClusteringIn order to address irregularities in the identified pixels of the images, the incorporation of clustering techniques was found to be necessary. The pixel identifications in the images exhibit a certain level of noise, as you can see in Figure 10.

The pixel identifications in the images exhibit a certain level of noise, as you can see in Figure 10. This is characterized by numerous small discontinuities and outliers.To tackle this challenge, we employed the DBSCAN (Density-Based Spatial Clustering of Applications with Noise) algorithm [7] for grouping the identified pixels into contiguous horizontal rows. DBSCAN is a powerful clustering method widely used in machine learning for separating clusters of high density from areas of low density, which, in our case, corresponds to distinguishing edge clusters from non-edge clusters.1635One of the key reasons behind our selection of DBSCAN was its robustness in handling noise and outliers. Designed specifically for noisy and sparse datasets, DBSCAN effectively differentiates between clusters and noise points by considering the density of the data points.

Designed specifically for noisy and sparse datasets, DBSCAN effectively differentiates between clusters and noise points by considering the density of the data points. Any outliers or isolated points that do not belong to any cluster are labeled as noise, enabling DBSCAN to maintain its reliability even in the presence of noisy data.An additional advantage of DBSCAN is its ability to automatically determine the number of clusters without requiring pre-specification. By dynamically detecting densely populated regions and ensuring their connection, DBSCAN offers a flexible approach for approximating layer boundary contours, which was highly beneficial for our research objectives.Figure 10 provides an illustrative example showcasing the application of DBSCAN to a sample ground truth image and a sample obtained through our simple thresholding method.

The effectiveness of DBSCAN in accurately clustering the pixels and identifying layer boundary contours can be observed.By utilizing DBSCAN in our analysis pipeline, we were able to overcome the challenges posed by the "messiness" of pixel identifications, enabling us to achieve more accurate and representative results in our research.Figure 10: DBSCAN clustering applied to a manually labeled image (top) and automated labeled image (bottom).Automated edge identification - Step 3: Function fitting1636We then implemented an interpolation function that drew a 1-D smoothing spline that fitevery group of edge cluster points of every image [6] (Figure 11).Figure 11: Interpolation function applied to pixel clusters.Performance evaluationThis step allowed us to compare and evaluate the error between the edges on each image labeled manually and its corresponding image automatically labeled. Performance metrics were computed to assess the effectiveness of each method.

Performance metrics were computed to assess the effectiveness of each method. To quantify the error between the sets of splines, we defined a function that takes two sets of splines as input. This function calculates the error by comparing the values of the splines at multiple points (in this case, 128 points) along the x-axis. The absolute difference between the spline values at each point is accumulated to compute a running total of the error.Subsequently, we iterated through pairs of ground truth and predicted sets of splines. For each pair, we calculated the error between the corresponding splines and tracked the total error for each set of splines. To determine the average pixel error for one pair of splines, we divided the total error by the number of pixels.Next, we calculated the average error for each set of splines (each image) by dividing the total average error by the number of splines in that set (image).

Additionally, we computed the average error for all sets of splines by dividing the total average error across all sets (images) by the total number of processed sets of splines.Finally, we calculated average pixel errors for all sets of splines. In cases where the number of predicted splines differed from the number of ground truth splines, we considered it a failure and reported failures separately.Results1637In this section, we present the results of our study on automated layer identification in 3D-printed walls from a LAAM printer. Figures 12 and 13, respectively, show examples leading to low and high accuracy for all methods. (The specific accuracy for each method depends on the light intensity variations over the image.) One can see the difference between the definition of the clusters for each automated method and our ground truth.

For the simple thresholding method, we observed that this method was sensitive to variations in image quality and noise, leading to inaccurate results. Next, we employed Otsu thresholding, which performed better than simple thresholding in terms of accuracy and was less sensitive to noise and image quality variations. However, it was computationally more expensive. Finally, Canny edge detection produced accurate results but was computationally most expensive and could be sensitive to noise.Figure 12: Example of Low accuracy DBSCAN clustering results on manually labeled image and automated labeled images1638Figure 13: Good Example of High accuracy DBSCAN clustering results on manually labeled image and automated labeled imagesFor each image, the three methods were applied in parallel, and the resulting layer identification outputs were compared against manual identifications.

We achieved highly accurate layer identification in 70% of the images using Otsu, 40% using Canny edge detection, and 30% using our simple thresholding method, with an average error of only 1 pixel per image when comparing the manually labeled and automatically labeled images for our best method. In Figure 14, we have the average pixel errors for all sets of splines. In cases where the number of predicted splines differed from the number of ground truth splines, we considered it a failure and factored it into the evaluation. This is shown in Figure 15.Figure 14: Average pixel error of edge identification1639Figure 15: Edge Identification Failure graphConclusion and Future WorkOur main goal is to develop an automated method for layer identification in 3D-printed walls produced by a LAAM printer. We tested three image thresholding methods: simple thresholding, Otsu thresholding, and Canny edge detection for layer boundary identification.

We tested three image thresholding methods: simple thresholding, Otsu thresholding, and Canny edge detection for layer boundary identification. Otsu thresholding was the most accurate method with minimal manual intervention, handling image quality variations and noise in approximately 70% of our images. Otsu thresholding is a promising approach for automated layer identification, providing high accuracy and reduced manual effort. Canny edge detection is effective but computationally complex. Simple thresholding may be suitable in specific scenarios, despite lower accuracy compared to the other methods.In the future, we would like to apply our methods to thermal images, to different types of materials, and potentially look into video streams. Finally, we expect to embed our approach to obtain an automatic image-to-layer data pipeline and integrate it into an optimization framework for manufacturing quality assessment and control.1640References1.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAn Automated Method for Geometrical Surface Characterization for FatigueAnalysis of Additive Manufactured PartsBehnam Rasoolian1*, Jonathan Pegues2,3, Nima Shamsaei2,3, Daniel Silva1,31Department of Industrial and Systems Engineering, Auburn University, Auburn, AL 36849 2Department of Mechanical Engineering, Auburn University, Auburn, AL 36849 3National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849Corresponding author:Tel: 334-439-9588Email: behnam@auburn.eduAbstractRecent interest to implement additive manufactured parts into structural applications has created a critical need to better understand the fatigue behavior of these parts.

Alloys such as Ti-6Al-4V are popular in the aerospace and biomedical industries due to their superior strength to weight ratio and biocompatibility. Previous works have associated fatigue behavior with surface roughness, especially radius of curvature of notches. It is therefore important to develop a fast, reliable and consistent methodology for extracting such curvatures. The contribution of this paper is in providing an automated method for extracting radius of curvature using image processing techniques and optimization. Results on fatigue life prediction indicates similar results between the automated method and manually extracted radii in a significantly shorter period of time.Keywords: Additive manufacturing, Fatigue behavior, Roughness, Radius of curvature2004IntroductionAdditive manufacturing has become an increasingly popular technique to produce structural parts or components.

In particular, additive manufacturing has gained a significant amount of interest in aerospace and biomedical sectors with Ti-6Al-4V alloy being prominently used in both these industries. The popularity of titanium alloys is due to their high strength to weight ratio, biocompatibility, excellent corrosion resistance, and general weldability. Studies have shown that variations in surface roughness can have significant effects on the fatigue resistance of laser beam powder bed fused (LB-PBF) metallic parts such as Ti-6Al-4V [1, 5].Pegues. et. al. [6] proposed an approach for estimating the effect of surface roughness on the fatigue strength of additive manufactured Ti-6Al-4V using an approximation of the fatigue notch factor, which was previously used to incorporate roughness in fatigue life in [7].

In order to completely analyse the effect of roughness on fatigue, these radii should be calculated for all roughness profiles. A typical surface roughness image can consist of hundreds to thousands of such profiles, making the measurements of these radii extremely cumbersome, if not impossible. An automated method is therefore desired not only to improve the time it takes to extract radii, but also to provide objective, reliable and consistent results that don’t depend on individual judgements.In this paper, we propose an objective, consistent, automated technique for extracting each radius of curvature of valleys as described in the proceeding section. The method significantly reduces the time it takes to find the radii, while providing results that are comparable to manually computed radii.

The method significantly reduces the time it takes to find the radii, while providing results that are comparable to manually computed radii. The proposed technique is aimed at reducing the time to manually measure the surface characteristics while improving the repeatability of the calculations to obtain accurate and consistent roughness measurements. These results can then be used as inputs for models that predict mechanical properties. Using an objective, repeatable method to calculate the inputs allows researchers to compare different models fairly.2005MethodsGiven a surface roughness profile containing point height data, obtained e.g. through microscopy, our method extracts radii of curvature for valleys. This is done by approximating the radius of curvature at the base of a valley which can be considered a local minimum in the roughness profile.

A 1-dimensional Gaussian kernel is used for smoothing. The size of the kernel is determined as a fraction (one fifth) of the average valley widths. The valley widths are determined as the distance between the beginning and the end of the valley that is below one standard deviation of the mean roughness. These local minima are then filtered such that only the lowest ones are chosen. The criterion for filtering is that a point is dropped if (cid:1877)(cid:3036) (cid:3408) (cid:1844)(cid:3028) (cid:3398) (cid:2026)(cid:3052) where (cid:2026)(cid:3052) (cid:3404) (cid:3495)(cid:2869) deviation of the profile. We denote (cid:1839) as the set of all valleys that are selected in this step. This process is illustrated in Fig. 1. The valley widths are depicted using black lines. The Gaussian smoothing kernel and the neighborhood from which the local minima are extracted both use kernels sizes that are one fifth of average valley widths.

This neighborhood ensures that a concave set of points are used to fit a circle. In order to guarantee that the neighborhood is symmetric relative to the local minima, the shorter distance from the two inflection points is chosen and mirrored to the other side of the valley (local minima). The neighborhood size normally does not exceed the size of the kernel used to find the local minima ((cid:1864)). In rare cases where this happens the neighborhood is shrunk to have size (cid:1864) ( on either side of the local minimum). Fig. 2 illustrates the (cid:2870) process of choosing the neighborhood. We let (cid:1840)(cid:3036) denote the points that comprise a neighborhood of the (cid:1861)-th valley in (cid:1839).(cid:3039)3- Roughness profile-Second derivativesInflection pointsChosen part of the curve to fit circle201I I I I I I I I II I I II:, . \ \ \ \ \... ·*The chosi n neighbo ood\ \ \\\ , ... ............... ......

\ \ \ \ \... ·*The chosi n neighbo ood\ \ \\\ , ... ............... ...... ......1.51.00.50.00.51.01.5Figure 2: Process of choosing a neighborhood for a typical valley. In order to maintain symmetry, the closest inflection point to the local minimum is chosen (in this case the inflection point that is to the right) and mirrored on the other side of the minimum.Finally, for each valley (cid:1861), a circle is fit to the points in (cid:1840)(cid:3036). In order to account for different scales (cid:4593). A circle is then fit to the and keep numerical stability, the points in (cid:1840)(cid:3036) are first normalized into (cid:1840)(cid:3036) normalized points and the resulting circle is transformed back to an ellipse in the scale of data points. Finally, the radius of curvature of the data points is computed by finding the radius of2007curvature at the base of the transformed ellipse.

To illustrate how the method works on a wide variety of profiles, radii extractions for several different profiles are provided in Appendix I.20081:e=;;g :t?:e-,;;;sa . I Et?S=:6:J :~so02575100125150175200225250275300325350400425450475500Length (µm)525550575800825850875900 Length (µm)925950Figure 3: Valley radii for a roughness profile. The golden horizontal line separates the valleys that are one standard deviation below (cid:1877)(cid:3364).Experimental ResultsThe proposed method of fitting circles was tested on local minima of randomly generated polynomial curves of degrees 2 and 3. For each polynomial degree, 1000 expressions were generated using coefficients that were randomly generated from values between 0 and 100.

For each polynomial degree, 1000 expressions were generated using coefficients that were randomly generated from values between 0 and 100. For example, to generate a single random polynomial of the third degree, 4 coefficients (cid:1853)(cid:481) (cid:1854)(cid:481) (cid:1855)(cid:481) and (cid:1856) were randomly generated and the expression (cid:1853)(cid:1876)(cid:2871) (cid:3397) (cid:1854)(cid:1876)(cid:2870) (cid:3397) (cid:1855)(cid:1876) (cid:3397) (cid:1856) evaluated. The radius of circles extracted was then compared to analytically computed radius of curvature at the local minima that where (cid:1876) is the location of the minima. The relative error were computed using the formula(cid:2869) (cid:3033)(cid:4594)(cid:4594)(cid:3118)(cid:4666)(cid:3051)(cid:4667)was then measured.

The relative error were computed using the formula(cid:2869) (cid:3033)(cid:4594)(cid:4594)(cid:3118)(cid:4666)(cid:3051)(cid:4667)was then measured. For polynomials of the third degree, the mean relative error was 0.0836 with standard deviation of 0.0893 and for polynomials of the second degree, mean relative error was 0.0166 with standard deviation of 0.0016. These results show that the fit circles are highly reliable measures of radius of curvature.The fatigue life estimations by Pegues et al. were reproduced using automatically extracted valley radii. In that paper, the authors provided fatigue life predictions of the endurance limit, using (cid:1837)(cid:4636)(cid:3033) (as defined in Equation 1), for two sets of roughness geometries, denoted AB1 and AB2. Using basic machine design principles, the authors then estimated the fatigue strength at 1000 cycles which has been traditionally used to signify the transition from low cycle to high cycle fatigue.

Using basic machine design principles, the authors then estimated the fatigue strength at 1000 cycles which has been traditionally used to signify the transition from low cycle to high cycle fatigue. A straight line fit in log-log coordinates is then constructed to estimate the stress-life curve. The important thing to note here is that only the smooth surface finish endurance limit, ultimate tensile strength, grain size, and roughness parameters are used to estimate the stress-life curves and that no fatigue data from the as-built surfaces are used in the predictions. Comparing the prediction curves with results from the destructive tests, it can be observed that the curve is slightly non- conservative.

Comparing the prediction curves with results from the destructive tests, it can be observed that the curve is slightly non- conservative. However, the predictions give reasonable results which, coupled with appropriate safety factors, may be useful in design given no available fatigue results are obtainable.A comparison of the predictions from the automated method and predictions of the original, manual method are demonstrated in Fig. 4. Graphs of the same color are predictions of fatigue life for parts with the same roughness geometries, with the dashed line corresponding to the manual method and solid lines corresponding to the automated method. Table 1 provides quantitative comparisons of the two methods in terms of mean squared errors (MSE) of the predicted fatigue life from the actual observations. As can be seen, the automated method can provide an accurate and reliable alternative for manually extracted radii in prediction of fatigue life.

As can be seen, the automated method can provide an accurate and reliable alternative for manually extracted radii in prediction of fatigue life. For both surface2009375975conditions, the automated method closely matches the results of the manual method. Furthermore, the automatic approach moves the predictive line closer to the fatigue data. The higher accuracy can be due to the fact that the automated method used a larger number of profiles to obtain valley radii. However, additional experiments are required to determine whether one method is consistently more accurate than the other. We can conclude that this approach can significantly reduce the time to obtain the radii measurement while also reducing the ambiguity inherent to manually measuring each radius.Figure 4: A comparison of fatigue life computations from Pegues et al.

The results are then used to estimate the effective stress concentration and fatigue notch factors for the AM surfaces, using the model proposed in [1]. Results show that fatigue life predictions using the automated method are comparable to results from manually extracted radii. The proposed extraction method can not only reduce the time costs required for manual measurements but, more importantly, also remove judgement element from the measurement, ensuring repeatability in the results despite the operator performing the experiments.2010References[1] J. Pegues, M. Roach, R. S. Williamson, and N. Shamsaei, “Surface roughness effects on the fatigue strength of additively manufactured Ti-6Al-4V,” Int. J. Fatigue, vol. 116, pp. 543– 552, 2018.[2] J. Gockel, L. Sheridan, B. Koerper, and B. Whip, “The influence of additive manufacturing processing parameters on surface roughness and fatigue life,” Int. J. Fatigue, vol. 124, pp. 380–388, 2019.[3] A. Yadollahi and N.

viewed PaperAutomated Path Planning for Wire Feeding in Large Format Polymer Additive ManufacturingMichael Borish*, Alex Roschli*Manufacturing Science Division, Oak Ridge National Laboratory, Knoxville, TN 37932AbstractPolymer-based large format industrial additive manufacturing (AM) technology continues to expandinto new application areas. One area of interest is large scale composite molds and dies. These molds and dies can be used for out-of-autoclave tooling applications. However, at these sizes, several challenges remain that prevent the use of AM technology due to cost. One such challenge is the need to heat these molds in large thermal ovens. To address this challenge, researchers at Oak Ridge National Laboratory developed the necessary hardware to allow co-extruded wire to be embedded into the material during construction. Using this hardware, a demonstration mold was successfully constructed and subjected to mechanical testing.

Using this hardware, a demonstration mold was successfully constructed and subjected to mechanical testing. The construction of this object required a unique pathing solution to achieve success. In this paper, we describe the needed software development in ORNL Slicer 2.0 to allow the automated production of this unique pathing solution.Keywords: slicing, additive manufacturing, wire coextrusion, wire embedding, self-heating molds, composite moldsIntroductionPolymer-based large format AM technology has proven attractive due to the large scale and speed with which objects can be constructed. The pellet fed systems typically use fiber reinforced polymer composites that have proven useful to multiple industries including aerospace, automotive, and renewable energy [1-3]. One such system shown in Figure 1 is Big Area Additive Manufacturing (BAAM) [4]. BAAM has successfully constructed molds for tooling applications [5].

One such system shown in Figure 1 is Big Area Additive Manufacturing (BAAM) [4]. BAAM has successfully constructed molds for tooling applications [5]. Large format polymer printers like BAAM have successfullyFigure 1: Big Area Additive Manufacturing (BAAM). Example large-scale gantry-based pellet fed polymer system.This manuscript has been authored by UT-Battelle, LLC, under contract DE-AC05-00OR22725 with the US Department of Energy (DOE). The US government retains and the publisher, by accepting the article for publication, acknowledges that the US government retains a nonexclusive, paid-up, irrevocable, worldwide license to publish or reproduce the published form of this manuscript, or allow others to do so, for US government purposes.

DOE will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).739demonstrated a significant reduction in the amount of time and cost of mold fabrication. These molds have been used in out-of-autoclave applications as an alternative to traditional autoclave approaches.Traditional autoclave approaches allow the construction of high-performance composite objects. Theprocess generally involves material laid upon a rigid mold that is then heated in an oven under pressure. Despite the success of this process, it is cost intensive in terms of capital, labor, and time [6]. To address some of these issues, out-of-autoclave (OOA) approaches were developed.

Despite the success of this process, it is cost intensive in terms of capital, labor, and time [6]. To address some of these issues, out-of-autoclave (OOA) approaches were developed. OOA generally involves vacuum-assisted consolidation with several distinct steps including: the construction of the mold, laying of material, void removal via vacuuming, curing in a thermal oven, post-cure heating, and removal [7]. OOA has been successful in general, but the OOA process still relies on the use of a thermal oven to cure the resin. Naturally, as molds scale in size, the cost of purchase and operation of such a large oven becomes an issue.One way to address limitations in the OOA process is the production of molds with embedded heatingelements. This construction is possible via coextrusion of a wire filament with a polymer additive process. Coextrusion is not limited to molds.

This construction is possible via coextrusion of a wire filament with a polymer additive process. Coextrusion is not limited to molds. There are numerous applications both small- and large-scale applications as well as various combinations of build process and coextruded material. Coextruded wire or liquid metal has been used for the fabrication of electronics and stretchable circuity [8-9]. Coextrusion is also not limited to a specific material type as continuous fiber for polymer composites, ceramics, and concretes has been investigated as well [10-14].Indeed, researchers at Oak Ridge National Laboratory have investigated coextrusion with the BAAMsystem in the past. While a full description of the hardware used is outside the scope of this paper, details can be found in previous work [15, 16]. An example of the hardware is shown in Figure 2.Figure 2: Co-extrusion hardware.

An example of the hardware is shown in Figure 2.Figure 2: Co-extrusion hardware. A custom wire feed system is attached to the nozzle of the BAAM system [15].In brief, the coextrusion system is comprised of a spool for material, a Midwest Motion Products DCmotor and gearbox, a custom wire feed system, a wire feeding tube, a wire-cutter with Bimba air-cylinder, and a dual-port nozzle. This hardware was utilized as proof of concept to construct self-heated molds on the BAAM system [17, 18]. While the proof of concept was successful and other examples abound, this work and others relied on a very manual pathing process. That is, pathing modifications necessary to correctly place the co- extruded wire were developed by hand, and this pathing was developed for relatively simple geometries.

That is, pathing modifications necessary to correctly place the co- extruded wire were developed by hand, and this pathing was developed for relatively simple geometries. Such a process is time consuming and is not practical for more complicated geometries.To address these manual limitations in the coextrusion process, researchers at Oak Ridge NationalLaboratory developed a methodology to construct molds with embedded heating elements automatically. To740produce such an object, a unique pathing solution was required. This work focuses on the development of a unique slicing approach integrated into ORNL Slicer 2.0. This algorithm allows the automatic generation of pathing necessary to construct these molds as well as features necessary for wire coextrusion.

This algorithm allows the automatic generation of pathing necessary to construct these molds as well as features necessary for wire coextrusion. In this paper, we outline these algorithms and provide a high-level overview of a produced demonstration piece and physical results.Algorithm Design and ImplementationTo provide an automated solution for the necessary pathing for wire feeding, ORNL Slicer 2.0 was modified with necessary functionality. To help visualize the goal of this algorithm, a representative rendering of the demonstration object with embedded wire is shown in Figure 3.Figure 3: Rendering of mold with embedded heating element [17].As can been seen from this figure, the mold with embedded wire is three beads in thickness. The wire embedded bead sits between the surface bead and the remaining geometry to ensure proper adhesion.

The wire embedded bead sits between the surface bead and the remaining geometry to ensure proper adhesion. For ease of discussion, the outer bead, wire embedded bead, and remaining geometry will be referred to as the surface region, wire region, and base region. Because the wire region must extend across the entire surface of the mold, pathing for the base region must be modified. Additionally, to ensure that the wire is taut during construction, sacrificial anchors are constructed on either side of this object. The purpose of the anchors is to provide a surface upon which the wire region can adhere as layer-wise construction continues. The wire region extends across the front of the mold to the outer boundary of each anchor.741To affect the necessary pathing solution, several significant components were developed in ORNL Slicer2.0.

These included settings mesh Boolean operations, anchor generation, unique base region pathing solution, and a restriction on post-processing travel optimization. The first feature that required expansion was the settings mesh. Slicer 2 allows a user to import multiple types of meshes. The most common mesh is the build mesh. This mesh represents the typical object in most slicers. Slicer 2 will cross-section this object and produce all necessary pathing based on user settings. In comparison, the settings mesh allows a user to modify existing pathing within a build mesh. An example of this is shown in Figure 4.Figure 4: Example use of settings regions to affect pathing in specific locations of build mesh.In this example, two settings meshes (transparent blue) were loaded on top of the square build mesh(green). In each of those settings meshes, the infill density was changed.

In each of those settings meshes, the infill density was changed. As a result, the area around the holes were densified to provide additional stock should the holes be bored. Settings meshes as presented in Figure 4 were limited to modifying pathing in a specific region. In this case, infill and its relative density. Settings meshes were extended to provide Boolean mesh operations such as union and intersection of 3D bodies. The Boolean mesh operations were provided as part of Computational Geometry Algorithms Library (CGAL) [19]. A full explanation of 3D mesh intersection is outside the scope of this paper. However, CGAL is one well- known computational geometry library that provides detailed explanations of provided algorithms. These Boolean operations are applied before the pathing of the object is computed.

These Boolean operations are applied before the pathing of the object is computed. An example is shown in Figure 5.Upper half of build mesh after Boolean operationArea contained within settings mesh after Boolean operationLower half of build mesh after Boolean operationFigure 5: Settings mesh performed Boolean intersection of mesh before pathing computation.In this case, the settings mesh (dark blue) performed a Boolean operation on the build mesh (light blue).Before cross-sectioning was applied, the settings mesh performed an intersection with the build mesh. The result from this intersection was the joined to the remaining geometry via mesh union. These geometric operations result three distinct islands for which pathing is computed. Because these regions have been split,742Slicer 2’s pathing algorithms evaluate them separately. These three islands are the area above, below, and inside the settings mesh.

Because these regions have been split,742Slicer 2’s pathing algorithms evaluate them separately. These three islands are the area above, below, and inside the settings mesh. This functionality was provided to allow a user to specify where in the object to embed wire along with other necessary user settings. Naturally, this process could also be affected by appropriate design via computer-aided design (CAD). If a user were to take the build mesh of Figure 5, split it into three separate STLs, and load those STLs into a slicing package, the same pathing could be produced. However, disassociating the various regions of interest into multiple independent STLs makes producing subsequent pathing more difficult. Rather, it is more straightforward to use the settings mesh to define an area of interest for wire coextrusion.For subsequent explanation, the setup in Figure 6A will be used as a reference.

Rather, it is more straightforward to use the settings mesh to define an area of interest for wire coextrusion.For subsequent explanation, the setup in Figure 6A will be used as a reference. This is a simple rectangular prism with a wire-embedded region specified by the settings mesh and represents the object shown in Figure 3. This setup is just before slicing has begun. With the settings mesh setup and appropriate user settings, slicing can begin. The first major step is cross-sectioning. As part of the cross-sectioning step, the build mesh was split according to the settings mesh. For this example, this process results in three areas of geometry: the surface region, wire region, and base region.

For this example, this process results in three areas of geometry: the surface region, wire region, and base region. A visualization of what the cross-section step processes is shown in Figure 6B.ABModified Cross-sectionUnmodified Cross-sectionObject BaseWire-embedded beadRemaining surface beadFigure 6:(A) Example mold (light blue) with settings region defining wire feed area (blue) just before slicing. (B) Cross-section output from first step in slicing process.During the cross-sectioning step, two sets of cross-sections are computed. One cross-section is of the original geometry and one is of the geometry after settings mesh Boolean operations have been applied. Both sets of cross-sections are carried through the steps of the slicing process.

Both sets of cross-sections are carried through the steps of the slicing process. The original cross-section will be used in combination with the modified cross-section to compute pathing for the base region of the object, while the modified cross-section will be used to produce pathing for the wire and surface regions. after the wire and surface region are removed and will be masked with the modified-cross section to manipulate the pathing. More details will be provided later in the section.With cross-sectioning complete, one final pre-processing step must be accomplished before pathingcomputation can begin: anchor generation. For the wire in the wire region to remain taut during construction, the wire must be anchored on both sides of the object. These anchors are automatically generated based on user settings during the slicing process much like other programmatic features such as rafts, brims, or skirts.

These anchors are automatically generated based on user settings during the slicing process much like other programmatic features such as rafts, brims, or skirts. These anchors are simple squares to which an additional wire region can be affixed to provide the necessary tension743during the construction process. Unlike the mold object, these anchors do not have an additional surface region since they will not be machined. Once construction is complete, the wires will be cut and the anchors discarded.Now that all appropriate geometry has been created, the slicing process can move to the path generation step. The pathing for the wire and surface region is computed using the modified-cross section. Both regions are typically one bead width wide. The pathing for these regions is straightforward and is the result of skeletons. Skeletons are open loop style paths that are typically used to fill in remaining voids inside of perimeters or other closed contours.

Skeletons are open loop style paths that are typically used to fill in remaining voids inside of perimeters or other closed contours. In this case, the skeleton pathing produces a simple straight line for both regions.No additional pathing is required for the surface region, however, the wire region requires one moreaddition. This region’s pathing must be augmented to include pathing for the attachment points at the anchor. Like pathing already produced, these anchor points are single bead skeletons appropriately distanced based on user settings. Importantly, this pathing is bundled with the wire region such that it is handled as a combined unit. This has the practical effect of enforcing a purely increasing or decreasing build order once travels have been added at a later step.Relatedly, the remaining pathing for the anchors is also computed. The anchors are programmatically generated based on user settings and are fundamentally like other features such as rafts.

The anchors are programmatically generated based on user settings and are fundamentally like other features such as rafts. Users specify general settings for size and gap distance for example and the structures are generated. These anchors consist of a single perimeter bead with an additional skeleton bead along the front surface where the wire will be placed. An example of this pathing is shown in Figure 7.Figure 7: Anchor example pathing. Pathing consists of single perimeter bead (blue) and single skeleton bead (red) for wire.The pathing for the base region is generated in parallel from the unmodified cross-section and is shownin Figure 8. Recall from the rendering in Figure 3, that the mold is three beads thick. The three beads are produced as a result of processing the unmodified cross-section as would happen in any normal slice. Once that pathing generation is complete, the outer two beads are altered.

The three beads are produced as a result of processing the unmodified cross-section as would happen in any normal slice. Once that pathing generation is complete, the outer two beads are altered. The outer two beads are intersected with the wire and surface regions from the modified cross-section. Polygon line-segment intersection can be decomposed into a simple line segment–line segment intersection algorithm. The Bentley-Ottmann algorithm is one such algorithm that allows computation of intersection points much faster than the naïve approach of evaluating every segment [20]. In brief, this algorithm is a line sweep algorithm that sweeps a line across all744segments that have been sorted by their x location for example. The resulting sweep produces a set of intersections that can be evaluated for path manipulation.Figure 8: Example mold with modified pathing.

The resulting sweep produces a set of intersections that can be evaluated for path manipulation.Figure 8: Example mold with modified pathing. Pathing consists of three modified perimeter beads (blue) and two skeletons (red): one for wire feed and one for the surface.Using these intersections, the beads are split with any pathing falling inside the wire or surface regions discarded. The modified cross-section acts as a mask to determine what pathing has not already been computed as part of the other regions. The remaining pathing is no longer a closed contour as a result, so a merge process is applied. Every pair of contours is spatially sorted to identify the closest points to connect. Once these points are determined, the pathing is connected to once again form closed contours. Using this approach, an arbitrary number of beads is possible with each odd multiple allowing two preceding outer-most beads to be connected.

Using this approach, an arbitrary number of beads is possible with each odd multiple allowing two preceding outer-most beads to be connected. In this case, the third bead is not altered as it does not intersect either of the other two regions.With all this pathing laid out, the final step is a post-processing step to add travels. Typically, thepathing is connected via travels according to user settings without restriction. In this case, some restrictions were imposed. The pathing that represents the wire region must be aligned and connected last, always. Additionally, the pathing for this region must always travel from the minimum to maximum or vice versa to guarantee the wire is stretched across the entire object and anchored appropriately. An example slice with numbered pathing is shown in Figure 9.235164 Figure 9: Example pathing of mold with anchors.

An example slice with numbered pathing is shown in Figure 9.235164 Figure 9: Example pathing of mold with anchors. Travels are inserted in numbered order to affect appropriate pathing.As shown in this figure, the travel process begins at the contour bead of the left anchor (1), travels to thebeads of the mold base (2-3), moves to the surface region (4), then moves to the right anchor (5), and finally constructs all three segments of the wire region (6). The travels insert with respect to all regions expect for the wire region can be arbitrarily decided upon, however, the wire region must always be done last and from one extreme to another. With this pathing in place, molds with embedded heating elements can be produced automatically.745Demonstration and DiscussionUtilizing the generalized approach described in the previous section, a demonstration part wasconstructed.

The part was constructed using the BAAM system with two different types of thermoplastic composite pellets. These two pellets were Polycarbonate with 20 wt.% carbon fiber loading (PC/CF) and polycarbonate with 20 wt.% glass fiber loading (PC/GF) (Techmer PM LLC (Clinton, TN, USA)). For the wire- embedded region, a nichrome alloy wire with Nickel 60%, Chromium 16%, and Iron 24%, and a diameter of 0.508 mm (McMaster-Carr, Cleveland, OH, USA) was used. The resulting object is shown in Figure 10.Figure 10: Demonstration mold with anchors and embedded heating elements [17].A full discussion of the mechanical properties of the mold can be found in previous work [17]. In summary, the mold was subjected to tensile, flexural, and thermal testing. Results showed that the mold displayed reasonable tensile and flexural behavior and was able to achieve a uniform thermal distribution.

Results showed that the mold displayed reasonable tensile and flexural behavior and was able to achieve a uniform thermal distribution. This is important as one of the primary motivations of this work is the high cost of large ovens in the OOA process. To be clear, these mechanical results and demonstration object were accomplished using a manual pathing process. That is, the pathing was manually created and laid out by hand. However, this process can now be created automatically with the implementation previously described.With respect to the algorithm presented in the previous section, the additional computation when compared to a typical slice is minor. Boolean mesh operations are accomplished via mesh refinement, and the operations are generally of linear complexity O(n). So, the cross-section step of the slicing process takes about twice as long. The only other significant aspect to the process is the manipulation of the pathing for the mold’s base region.

So, the cross-section step of the slicing process takes about twice as long. The only other significant aspect to the process is the manipulation of the pathing for the mold’s base region. The Bentley–Ottmann algorithm processes a sequence of 2n + k events, where n denotes the number of input line segments and k denotes the number of crossings. Each event requires log n time to evaluate. Ignoring constants, the complexity is thus O ((n + k) log n). However, nearly all of the computational infrastructure is already in place due to the typical slicing process. Practically, this means that the additional computation adds only a few seconds to the typical slice.Computational efficiency aside, the other practical benefit of this approach is the ability to support morecomplex geometries. While the implementation and examples to this point have been simple rectilinear geometries, this has been for ease of demonstration.

While the implementation and examples to this point have been simple rectilinear geometries, this has been for ease of demonstration. There is nothing in this approach that requires rectilinear geometry. The algorithm as described is capable of ingesting meshes of any shape and complexity and producing pathing with the necessary commands for wire coextrusion. The only requirement is that a designer be able to produce a settings mesh representative of the volume in which wire coextrusion is expected to occur in.746ConclusionAs can be seen, the proposed algorithm requires only a minimal increase in computation to provide anautomated solution. This pathing solution has been shown to allow the production of molds with embedded heating elements. Through physical testing, these molds were found to have appropriate tensile and flexural characteristics as well as uniform heating properties making them suitable for OOA applications.

Through physical testing, these molds were found to have appropriate tensile and flexural characteristics as well as uniform heating properties making them suitable for OOA applications. This construction was made possible by a wire coextrusion system attached to BAAM, a large format polymer platform. Though the presented approach has a number of advantages, there is still room for improvement. Future work will focus on the automated identification of the most appropriate region for wire coextrusion. Currently, this process requires the use of a second mesh, however, its possible that this need not be the case.AcknowledgmentsThis material is based upon work supported by the U.S. Department of Energy, Office of EnergyEfficiency and Renewable Energy, Office of Advanced Manufacturing, under contract number DE-AC05- 00OR22725.BibliographyK. Friedrich, A.A. Almajid, Manufacturing aspects of advanced polymer composites for automotive applications, Appl. Compos. Mater.

tomated Planning for Material Shaping Operations in Additive/Subtractive Solid Freeform FabricationJianpeng Dong, Ju-Hsien Kao, Jose M. Pinilla, Yu-Chi Chang and Fritz B. PrinzStanford University Department of Mechanical Engineering Stanford, CA 94305-3030Abstractremoval processes, additive/subtractive solid freeform fabrication (A/S SFF) can build parts with complex shapes without compromising precision requirements. However, preparing material removal operations requires special expertise, which has in fact become one of the bottlenecks of the A/S SFF manufacturing process. To achieve automated planning, a shaping process planner is being developed based on. 3D solid representation and a surface classification scheme. This planner can generate numeric control (NC) codes for CNC milling in an automatic fashion on non-undercut features of arbitrary 3D input geometry. Planning approaches are also presentedinorde~to>shape parts accurately and efficiently.

Planning approaches are also presentedinorde~to>shape parts accurately and efficiently. The proposed shaping planner thus delivers on the promise offully automated process planning in A/S SFF.Combining the advantages oflayered manufacturing and materialKeywords: CAD, CAM, Solid Freeform Fabrication, CNC Machining, Process Planning, Tool Path Generation, Automation1. IntroductionIn most Solid Freeform Fabrication (SFF) .. syste11ls, CAD models are first decomposed into 2D layers, and then physical •.· parts are built up layer by layer. Decomposing a 3D model and building a physical part in this fashion, however, can result in stair-steps on freeform surfaces; thus sacrificing surface finish and dimensional accuracy of the resulting parts [2].the advantages of layered manufacturing and material removal processes, such· as .CNC 111illing, so it can build parts with complex shapeswithoutcompro111isingprecision requirements. However, integration of material removal.

However, integration of material removal. operations with material. addition operations makes. process planning. for A/S SFF much more challenging than that .for purely. additive SFF [1]. In particular, preparing for shaping operations and generating numenc control (NC) codes ·for CNC milling.require<special expertise. and. experience, which have in fact become one of the bottlenecks of theA/S SFFmanufacturingprocess. To make A/S SFF more productive, a process planner for shaping operations is indispensable.Additive/subtractive Solid Freeform Fabrication (A/S .SFF)combinesIn this paper, A/S· SFF. process will be examined .from a process planning perspective. Then how to automate shaping operations will be considered. Based on a proposed. surface .... classification scheme, strategies are .presented to improve machining efficiency and maintain part accuracy. Finally, an example part is analyzed to show the capabilities of the proposed planner.1212.

Finally, an example part is analyzed to show the capabilities of the proposed planner.1212. Shaping Process PlannerNS SF,F processes [1-31 involv~ iteration of material deposition and·· material removal operatIons, as can be seen from FIgure 1. Starting with build direction selection it then deco~n:poses a c;AD mo~el into 3D layers with variable. thickness using a.3D model decOmpOSItIon technIque. ThIs technique can identify parting lines of undercut and non undercut features ota model, and uses these parting lines to splitthe model [2, 4]. Part and support structures· are thus constructed with extrusion operations in the computer. As a result, undercut features of the model will become non-undercut features of one of the decompose~ 3D lay.ers; thus no undercut.

AlS SFF process planningAs discussed above, a distinct characteristic of shaping operations in AlS SFF· is that only non-undercutfeatures. of the decomposed 3D layers need to be machined during manufacturing.• Decomposing CAD models in this way also avoids fixture design. for shaping these 3D layers·because previously builtlayers form the natural support structures. In addition,model decomposition.takesremovalprocess constraints,.such .as cutter length, into account; thus making all non-undercut surfaces· of each 3D layer accessible by cutters from above. Therefore, shaping process planning becomes easier than conventional CNC machining.An automated shaping planner has been designed to be capable of generating NC codes for any given 3D layer without human intervention, which is shown in Figure 2. The inputs for this planneraredecornposed3Dlayers, and the outputs are the NC codes for shaping the layers.• It alsoineeds •• the cutter geometry information before tool.

The inputs for this planneraredecornposed3Dlayers, and the outputs are the NC codes for shaping the layers.• It alsoineeds •• the cutter geometry information before tool. paths can be generated and machine information before post-processing .•• can be .• accomplished. The implementation of lies on top of commercial CAD/CAM packages, Unigraphics, which provides toolkits to. access to CAD model database and generate. tool paths.this. ·.plannerIn this paper, the discussion will focus on planning fin~s~ing operations .o~ 3-~xis CNC milling machines because finishing operations take the maJonty of the machirnng tune and are crucial to part dimensional accuracy and surface finish.Figure 2. Shaping planner3. Tool Path GenerationThe essential tasks of a shaping process planner are to select and organize different shaping operations andto generate toolcutting paths over the model surfaces to be shaped.

Tool Path GenerationThe essential tasks of a shaping process planner are to select and organize different shaping operations andto generate toolcutting paths over the model surfaces to be shaped. These selection•••and organizationdecisions are based·· on \Vhattool path generation methods are available and how tool paths are generated. A variety oftool path generation methods have been developed. Most of recentresearch work focuses on generating tool paths over freeform surfaces. Usually, a model is a mixture of a variety of surfaces; hence an automated planner should deal with the whole .model instead of individual surfaces. But generating tool paths over different types of surfaces is rarely studied. Several non commercial approaches have been proposed,but they are either impractical or too hard to have a robust implementation [5-7, 10].In this study, two widely available types of tool pa.ths, layer-based and projection based, are used.

In the case of layer-based methods, •a grol.1pofparallel planes of constant (which is called the. cutting depth) are used to slicethe\Vhole model,and spacing intersection. curves of these planes with the model are offset with respect to the. cutter geometry to form the tool paths. In the case of projection..basedmethoqs,alooPiofcurves are selected as boundary curves. The boundary curves are. then recursivelYi()~fset by a constant distance (which is called. the st~p-overdistance).• The resulted offset curves are projected ontothe up-facingsurfaces in order to update the z coordinate of each pointof the offset curves.. These modified curves are the tool paths.Th~ .most imp0rtantparameters controlling the 111achinillgtolerance arecu1ting .depth (in the case oflayer-based tool paths) and step-over distance (in thecaseofproj~ction­ basedtool paths).• Both ofthem/are required. by •. theQ~i~apkages ·.. tobecQllsblllt idllring tool pathgenera.tion.

by •. theQ~i~apkages ·.. tobecQllsblllt idllring tool pathgenera.tion. Some variatiQnsdoexist[8,91.Forinstance, the whole m()d~Lcanbe divided into several blocks. Within each block, the constantconstraint is required, but thevalues chosen in different blocks may be different. In addition, the number of such blocks are limited. Owing to these constraints, each method may produce longer tool paths than necessary to meet the machining tolerance. If the cusp height of material left over after the model is machined is used as a measurement of the machining tolerance, how the ideal cutting depth or step-over distance changes as the normal of a to-be-shaped surface changes can be shown in Figure 3.Figure 3 (a) shows an inclined face being shaped with a ball endmill.

The relationship between machining tolerance 8, cutter radius r, face inclining angle a, step over distance s and cutting depth c is given in the follows expressions:c = ~r2 - (r - 8)2 sin a; s = ~r2 - (r - 8)2 cos a;(1)(2)the ideal cutting depth of layer-based tool paths will have the minimum value when the inclining angle a is equal to 0°, and it will have the maximum value when a = 90°; On the other nand, the ideal step-over distance of projection-based tool paths will have the minimum value when a = 90°, and the maximum value when a =0°. It can also be seen that when a =45°, the ideal cutting depth and step over distance have the same value. Owing to the restriction that both cutting depth and step over distance should be keptconstant during tool path generation, the smallest cutting depth or step·over distance value has to be chosen. when a freeform surface is to be shaped; otherwise,.

Normal direction effectoncutting depth and step-over distance. Cutter radius is 12.7 mm, and cusp height is 0.003 mm.904. Surface ClassificationConsidering the effect of surface steepness, non-undercut surfaces (because in A/S SFF only non-undercut surfaces of a compact are to be shaped) are classified into three (2) categories: horizontal faces which have normal parallel to the z direction; (3) freeform surfaces or inclined faces. Furthermore, freeform surfaces and inclined surfaces are identified as flat and steep surfaces. The criterion for distinguishing a flat surface from a steep one is to use a 45° angle between the surface normal and the build direction. If the surface normal forms an angle less than or equal to 45° with the z direction, this surface will be identified as steep surface; otherwise, it is flat.

If the surface normal forms an angle less than or equal to 45° with the z direction, this surface will be identified as steep surface; otherwise, it is flat. If necessary, a freeform surface will be split into flat and steep portions based on the same criterion.(1) vertical walls which have normal perpendicular to the z direction;Different types of surfaces should be shaped differently. For instance, to machine vertical walls or horizontal planar surfaces, flat end-mills are usually used; while ball end mills are usually used to machine curved freeformsurfaces or inclined faces on 3-axis CNC machines. To enhance machining efficiency, freeform surfaces or inclined faces are further categorized based on their surface normal values. The details will be illustrated in Section 5.S.Machining Strategies5.1 Hybrid-machiningMachining using either layer-based or projection-based approaches alone is inefficient because the cutting depth or step-over is restricted to be constant.

One solution is to apply layer-based and projection-based tool paths one after the other, shaping the whole model. In this case, the smallest ideal step-over distance or cutting depth value does not have to be used. Instead, the corresponding values at a 45° angle are used. The reasons are as follows. If the ideal step-over distance at 45° can achieve the specified machining tolerance, then using the same step-over distance will produce smaller cusp heights over those surfaces with a normal greater than 45°. Similarly, if the cutting depth at 45° can achieve the specified machining tolerance, then using the same cutting depth will produce smaller cusp heights over those surfaces with a normal less than 45°.

By applying both tool paths, the whole model is guaranteed to be shaped to the specified machining tolerance.Figure 4.....'I..'"'~u.,,~vJ.J.,:u."u~Jt::J between different machining approaches125This strategy significantly reduces the machining time even though almost every As can be surface .will have to. be shapecitwice.A.siITlpleexample is shown. in Figure seen,aihalf-spherewitharadiusQf.50nnn is. being shaped> by· a ball enci cutter with a radius of 12.7mm. The cutting feed rate is 1000.0 mmlmin. Nocu.tterchange or tool e~gage andretract time is considered. If each type of tool paths is appliedalone,Joshape thIS model to the specified tolerance, which is assumed to be 0.003 111m, more than 20 hours h,as tobe.sIJent onmachil{ing,ascanipeseen froll1 Table 1. ByapplY1:ngpoth tool paths, sInce a greater step-over dIstance and cutting depth can be used, the machining time IS. reducedby two orders ofmagnitude.

ByapplY1:ngpoth tool paths, sInce a greater step-over dIstance and cutting depth can be used, the machining time IS. reducedby two orders ofmagnitude. Obviously, this· strategy is still not efficient because almost every surface of the model has been machined twice..1. ...J..~~\"ILlJ..l.LUJL;:;., efficiency comparison of different tool path generation approaches5.2 Surface-splittingTo avoid machining every surface twice, a surface splitting approach is proposed to distinguish surfaces based on surface normal distribution. First, classify all the surfaces based on the scheme presented above. If necessary, split freeform surfaces and inclined surfaces into flat and steep portions..

First, classify all the surfaces based on the scheme presented above. If necessary, split freeform surfaces and inclined surfaces into flat and steep portions.. Generate tool paths on vertical walls or horizontal planar faces usingflatend,-mills, and generate layer-based tool paths on steep portions and projection-based tool paths on flat portions separately using ball end-mills.A similar algorithm to model decomposition has been developed in .order to find the boundary curves between flat surface portions and steep ones in the model. In this case, instead of finding the parting lines (each point on which has a normal perpendicular to z direction), the splitting curves (each point on which has a normal parallel to a planar face with either a.45° inclination angle) need to be identified. Once the splitting curves are found, they are used to identify which.

Once the splitting curves are found, they are used to identify which. portions of a surface should be machined using layer-based tool paths and which portions should be shaped using projection-based tool paths. Furthermore, same type surfaces can be grouped together; thus reducing the total number of operations.it reduces machining time by about 1/3 compared with the above hybrid-machining method, as can be seen in Table 1.If this approach is applied to the same model shown in Figure 4,6. Case StudyThe proposed shaping planner can be used to shapenon-updercut features of any 3D layers and output theNC codes with few i limited manual inputs, such as cutter information (tool number) and machining tolerance. Currently, it also needs to manually identify narrow deep cavities and select.

Currently, it also needs to manually identify narrow deep cavities and select. different cutters based on .the geometry of such cavities, But an automated cutter selection modtllewillbedeveloped as a part of the planner based on a medial axis transformi(MAT) representation of the model [11].Figure 5 shows two sample parts built with this planner. The right photogr~ph shows two parts shaped with different approaches. The above one was shaped only uSIng projection-based tool paths. The bottom one was shaped using hybrid-machining approach. the Both of them use the. same step-over distance. As can be seen from this photograph, hybrid-machining approach exhibits better surface quality. Because a greater .step-over distance and cutting depth were adopted, the machining efficiency was also enhanced.(a)(b)Figure 5. Sample parts built using different approaches.

Because a greater .step-over distance and cutting depth were adopted, the machining efficiency was also enhanced.(a)(b)Figure 5. Sample parts built using different approaches. (a) Sample parts, (b) parts shaped by projection-based tool paths alone (above) and by hybrid-machining approach (bottom)7. ConclusionsA generic shaping operation planner purely based on model geometry is presented. Using widely available tool path generation methods, its implementation is>simple and its performance is .•• robust. Two approaches, hybrid-machining and. surface-splitting, are proposed .to reduce machining time and enhance efficiency of the material removal operations in AlS SFF.• In particular, the proposed surface-splitting approach can reduce machining time by two orders of magnitude. This method splits the surfaces into different portions or distinguish individual surfaces into different types based on surface normal.

This method splits the surfaces into different portions or distinguish individual surfaces into different types based on surface normal. Tool paths are then generated by altemativelY.iapplying.layer-based and .projection-based methods on different types or portions of surfaces, This planner will not only reduce the planning overhead introduced by the shaping operations in AlS SFF, but it will eventually become part of a fully automated AlS SFF process planner.AcknowledgmentsThe authors acknowledge support from the Defense Advanced Research ProjectsAgency, the Office of Naval Research andUnigraphics Solutions Inc.R.eferences[llJ.M.iPiniHa, J.,.I(ao andF. .B..Prinz,i1998, ..• Pro.cess .Planning and.Automationfor Ada.itive..SubtractiveSolidFreeform Fabrication, SFF Symposium 1998,Austin,·TX[2lL.Weiss,R. Merz, et aI, 1997, Shape Deposition Manufacturing of Heterogeneous Structures, Journal of Manufacturing Systems, 16 (4), 239-48127[3] A. G. Cooper, S.Kang, J. W.

tomatic CAD-model Repair: Shell-ClosureJan Helge B~hn and Michael J. WoznyRensselaer Design Research Center Rensselaer. Polytechnic Institute Troy, New York 12180-3590 (518) 276-6751 bahn@rdrc.rpLeduAbstractShell-closure is critical to the repair of CAD-models described in the .STL file-format, the de facto solid freeform fabrication industry-standard. Polyhedral CAD-models that do not exhibit shell-closure, i.e. have cracks, holes, or gaps, do not constitute valid solids and frequently cause problems during fabrication. This paper describes a solution for achieving shell-closure of polyhedral CAD-models. The solution accommodates non manifold conditions, and guarantees orientable shells.

This paper describes a solution for achieving shell-closure of polyhedral CAD-models. The solution accommodates non manifold conditions, and guarantees orientable shells. There are several topologically ambiguous situations that might arise during the shell-closure process, and the solution applies intuitively pleasing heuristics in these cases.1 IntroductionIt is critical that the CAD-models used in solid freeform fabrication (SFF) are robust to avoid failure during manufacturing. Unfortunately, CAD-models described in the .STL file-format [1], the de facto industry standard, frequently lack this robustness [2, 3, 4]. Typical problems include punctured shells, inconsistent facet-orientations, and internal walls and structures.the CAD-model solid is defined by its surface, Le. the If this shell is punctured (i.e. solid is defined by a closed shell enclosing its material.

the If this shell is punctured (i.e. solid is defined by a closed shell enclosing its material. open), then it can no longer properly contain the material, and this consequently causes the failures commonly associated with cracks, gaps, and holes (Figure 1.1). Similar problems arise when the facet-orientations across a shell are inconsistent. This too causes uncertainty regarding what side of the shell contains material. Finally, structural problems might arise from internal walls and structures within the solid. These internal features can create irregularities in the solidified material during the processing. For instance, with stereolithography [5] and selective laser sintering (SLS) [6], excess amounts of unsolidified material might get trapped inside the solidified matrix, and consequently result in uncontrolled shrinkage and surface warping (Figure 1.2).With the .STL file-format,The .STL file-format is capable of representing a robust, well-formed solid.

Unfor tunately, contemporary CAD-systems frequently fail to achieve this goal. There are twoCopyright © 1992 Rensselaer Design Research Center. All rights reserved.86Slice Plane..Stray Scan-VectorsFigure 1.1 Shell-punctures causes the laser to produce stray scan-vectors.Figure 1.2 Internal walls cause structural discontinuities that can lead to excessive shrinkage.design environments currently available for CAD; solid- and surface-modelers. Solid modelers generally avoid punctured shells, but frequently generate needless internal walls In addition to these problems, surface and structures, and inconsistent orientations. modelers frequently fail to properly mate adjacent surfaces; this results in punctured shells.These problems are commonly resolved by the SFF-equipment operators prior to fabrication. The operator might attempt to edit the .STL file directly, or, if available, edit the original CAD-model to eliminate the problems.

The operator might attempt to edit the .STL file directly, or, if available, edit the original CAD-model to eliminate the problems. However, manual model repair becomes excessively expensive when the number of facets reaches 10,000 to 100,000+, and the number of errors run into the hundreds. This has lead to the development of automatic CAD-model repair.Brock Rooney & Associates [7] perform limited CAD-model repair in their .STL translators, ensuring that their .STL output does not contain punctured shells. Their pro prietary shell-closure solution is not complete. It does not handle orientability-problems satisfactorily, and it has limited success with certain simple structures. However, they report few problems in general with their .STL output when run on the software-robust stereolithography process of 3D Systems, Inc.

However, they report few problems in general with their .STL output when run on the software-robust stereolithography process of 3D Systems, Inc. In the rare cases where their shell-closure algorithms fails, the CAD-model must either be regenerated at the design-level, or be repaired manually as described above.3D Systems, Inc. [8] provides its customers with a limited, proprietary model-repair facility. It too is intended to ensure that the model does not fail during fabrication, or if the repair fails, to provide indications to where it must be repaired manually. The program attempts to repair simple orientation problems, minor punctures, and facet intersections, but leaves internal walls, and fails for large punctures and non-orientable shells. It performs the repair under the restriction of several user-specified parameters, and if these are set too conservatively, an intermediate result is generated.

It performs the repair under the restriction of several user-specified parameters, and if these are set too conservatively, an intermediate result is generated. This output can be processed iteratively until a valid, though not necessarily correct result occurs.The basic problem of CAD-model repair is that of inferring information that has been lost somewhere in the design process. While it might be easy in some cases to detect the presence of a problem, it is frequently impossible to determine with certainty what the correct solution is. This is especially true if the designer's intent is lost. Without it, an educated guess must be made whenever an ambiguity is encountered.

This is especially true if the designer's intent is lost. Without it, an educated guess must be made whenever an ambiguity is encountered. The goal of our work is to match the quality of manual repair that is performed without the knowledge of the designer's intent.87\Figure 2.1 A non-manifold condition can be viewed locally as a set of two-manifold conditions.2 Shell-ClosureIn This paper presents a complete solution for shell-closure of polyhedral shells. doing so, it also solves the orientability problems, and it simplifies the complexity of the model-topology for more efficient processing during fabrication. In particular, it ensures a pseudo two-manifold topological organization of the facets to facilitate high-speed slicing [4, 9].

In particular, it ensures a pseudo two-manifold topological organization of the facets to facilitate high-speed slicing [4, 9]. This topological organization also simplifies the subsequent resolution of shell intersection and shell-nesting problems, which are necessary for the removal of needless internal walls.A (two-) manifold structure is homeomorphic to a two-dimensional disk at every point across its surface, i.e., each such point has a neighborhood that is topologically equivalent to a two-dimensional disk. A non-manifold structure violates this constraint; two cubes that touch along a common edge is one such example. However, if the facets at a non-manifold location are paired up properly, then this situation can be described as a set of (pseudo) two-manifold conditions [10] (Figure 2.1).Our solution for shell-closure follows a three-stage strategy: First, we identify the lamina edges.

Second, we combined these edges into closed, non-self-intersecting loops of boundary-curves; commonly known as directed Jordan curves. And finally, we create lids that fit these Jordan curves and that close the open shells.3 Lamina Edgesit A lamina edge is generally considered a special case of a non-manifold edge; has only a single adjacent face [11]. Lamina edges are of interest because they make up the boundaries of the shell-punctures, and of the corresponding lids which we will create to close the shells. However, if the definition of a lamina edge is generalized, in particular generalizing the definition of a puncture, then we can expand the use of creating lids for shell-closure to also solve erroneous facet-orientations and complicated non-manifold conditions.Expanded definition: A lamina edge exists along each face-edge that does not mate and pair properly this particular face with any other face along this edge.

These lamina edges can be found as follows: Let all face-edges be candidates for lamina edges. Then, for each face-edge, attempt to find an adjacent face of compatible orientation such that883.1 Two laminaof two faces aboutSurface Normal}={... }candidates (a,b) are removed from the set of candidates (LEe) after a successful pairing3.3 Two unpaired facets along a non-manifoldyield two coincident lamina edges.adjacent about the common edge. are removed from the set of lamina after this process is exhausted,4.... WOIlA.'Vl"LLLj;.

are removed from the set of lamina after this process is exhausted,4.... WOIlA.'Vl"LLLj;. ULL.v'VIL,VU.curves from the set of lamina edges, .UAl..'V "'v"'"',,...u.F-, boundary curve-segments, or to a single shell, and is terminated'"''VA.'''''Figure 4.1 Simple, self-connecting loop (unambiguous).Figure 4.2 Linear combination (unambiguous).Figure 4.3 Cyclic with single direct returns.Figure 4.4 Cyclic with multiple direct returns.Figure 4.5 Trivial and nontrivial cases of cyclic with non-direct returns.by two endpoints that might coincide. This can be achieved as follows: Initially, consider each lamina edge as an arc. Then, iteratively, for each vertex with only two adjacent arcs, merge those two arcs into a single arc, causing this vertex to no longer have an adjacent arc.

Then, iteratively, for each vertex with only two adjacent arcs, merge those two arcs into a single arc, causing this vertex to no longer have an adjacent arc. The final set of arcs form a set of directed graphs (as defined in graph theory [12]) of various complexities that describe the shell-puncture boundaries, and these puncture-boundaries must be separated prior to lid-creation.We have identified five classes of puncture-boundaries of increasing identification complexity, which when combined, form the above mentioned sets of directed graphs. Among these classes, three occur in our industrial .STL test-suite, while the remaining two, the most complex classes, persist as unobserved possibilities: The most trivial class of all is the simple combination; a single self-connecting arc.

It is unambiguous and trivial to detect, and it forms a complete puncture-boundary by itself (Figure 4.1).The linear combination is also unambiguous; two arcs that are adjacent to the same shell and that form a closed loop. These arcs form a link between two separate subgraphs, each of which is less complex that their combination (Figure 4.2).The remaining three classes are ambiguous because they offer multiple closed-loop alternatives, and because the Jordan curves are located in three-dimensional space rather than in a simplifying two-dimensional plane. Fortunately, one can make some reasonable,90aalternatives: For instance, separate shells. This is based solids; consequently, all facets shell. This dictates that, for same shellnot be adjacent to....V ...'VA-A-LJV'Y by the formation of cyclic i"""Tl'~i"n1nn arc for each of its has been eliminated as our rare, but present,not yet en- ambiguous problem arc, multiple options remain.

It U ........ILLLd'AAE:, not only topological-, 'l"nQTf't"I1nn problem where the connecting .......u.J--,v.u ...u of case, and it has a graceful more ambiguous. Fur with minimum weight problem arc-pairs, we end up with Jl"'-V ..........JlAA..., back to the example in arcs are not shown), yielding arc), the above shell-merger downward arc) as an ac- V"'''''.l.AAJUA..., arcs as possible matches will cause the merger of two where further loop-extraction two possible combinations....AA'L,' .... "'.AA wereturns is anthese are combinations ..... .I.'V''-'''''' ..... loop (Figure 4.5). Some .I.......,.l.V,....."" .. A ........I..I. 7 ........'.I....,JL......., arcs into longer parts, but and ambiguous. One possible are more arcs onto a suitable two-manifold structure,...... ""'.... from the more complex these problems appear to our strategy is to identify the less original graph. This simplifies ......... ' ........... 'VAA ....... unambiguous constructions. AsV./l........

from the more complex these problems appear to our strategy is to identify the less original graph. This simplifies ......... ' ........... 'VAA ....... unambiguous constructions. AsV./l........ ""'..,sub-problems.orr([r 3 (f'r125674r r1'89Figure 5.1 Lid-creation by iterative removal of the vertex with the smallest angle.Figure 5.2 Iterative post-processing to improve the quality of the lid-surface.5 Creating LidsThe task of creating lids is greatly simplified once the corresponding directed Jordan curves have been identified: The puncture is closed by an iterative algorithm that decrements the Jordan curve by eliminating one vertex at a time; the vertex is eliminated from further consideration by adding a facet that spans it and its two neighboring vertices along the shrinking Jordan curve.

At each iteration, we remove the vertex along the Jordan curve that spans the smallest angle with its two neighboring vertices along the remaining, decremented Jordan curve (Figure 5.1). This avoids the clearly undesirable splitting of angles, and it will in general produce excellent results, especially for long narrow punctures which abound in industrial .STL files [7, 8].It is a time-local algorithm in that it finds the best selection at any given time, without keeping the overall optimal solution in mind. An example of this is the selection of the third facet in Figure 5.1; it forces the subsequent undesirable creation of facet number four.

An example of this is the selection of the third facet in Figure 5.1; it forces the subsequent undesirable creation of facet number four. Fortunately, such problems are easily rectified with a simple, inexpensive iterative post processing step as illustrated in Figure 5.2: For each adjacent pair of lid-facets, flip their common diagonal edge if this increases their minimum facet-angle.However,this heuristic algorithm is not without drawbacks.A more serious problem is the difficulty of finding the angles spanned about the vertices. While angles between vectors in 3D-space are in the range 0-180 degrees, it is clear, from the 2D-plane, that the angles between facet-edges are in the range 0-360 degrees. It turns out, however, that a conservative decision to assign the range 180-360 degrees to a vertex whenever there is a doubt to which half of the total range it belongs to, simply postpones its elimination.

Eventually, one of the vertex's neighboring vertices92Orientation DiscontinuityFigure 5.3 Cutting a non-orientable Mobius-strip along its orientation-discontinuity, results in a orientable panel.will be eliminated, and its angles must be recomputed. Usually, this clarifies the situation, and the vertex might be assigned the range 0-180 degrees instead.It is interesting to note that the above lid-creation algorithm is capable of creating a non-orientable shell if the open shell is non-orientable. An example of this would be the closure of a Mobius-strip. However, with our definition of lamina edges, non-orientable In particular, the shells are cut such that they become orientable prior to lid-creation. Mobius-strip becomes a simple panel (Figure 5.3).

Mobius-strip becomes a simple panel (Figure 5.3). As a consequence, since the lid creation algorithm maintains the directionality of the diminishing Jordan curves, we are guaranteed to maintain orientability throughout the shell-closure; and more importantly, the final closed shells are therefore always orientable.6 ConclusionsThis paper has presented a complete solution for shell-closure of polyhedral shells, the first step towards well-formed and manufacturable solids. The solution is based on topological principals, and consequently resolves orientability and non-manifold problems in addition to eliminating shell-punctures (i.e. cracks, holes, and gaps).

cracks, holes, and gaps). As a result, the facet-model output consists entirely of closed, orientable shells, and the resulting facet models can therefore be viewed as a set of two-manifold shells.This topological organization simplifies the subsequent resolution of shell-intersection and shell-nesting problems, and it is a prerequisite for the detection and removal of zero volume parts and other needless internal walls that create structural discontinuities in the fabricated parts.7 AcknowledgmentsThis paper is taken in part from a thesis to be submitted in partial fulfillment for the degree of Doctor of Philosophy in the Department of Electrical, Computer, and Systems Engineering at Rensselaer Polytechnic Institute.This research was supported by NSF Grant DDM-8914212 as a subcontract through the University of Texas Solid Freeform Fabrication program, and other grants of the Rensselaer Design Research (RDRC) Industrial Associates Program.

Any opinions, findings, conclusions, or recommendations expressed in this publication are those of93the authors and do not necessarily reflect the views of the National Science Foundation or any of the industrial sponsors.We would like to thank Dick Aubin, Pratt & Whitney (a division of United Tech nologies), for providing a number of industrial .STL models. Also, special thanks to the members of the RDRC: Stephen Rock for his early work on this project and his subse quent collaboration; James Miller and Robert O'Bara for their development of facet-model visualization tools; and Peter Wilson for his helpful review of this paper.References[1]"Stereolithography Interface Specification," 3D Systems, Inc., June 1988.[2]S. 1. Rock and M. J. Wozny, "A Flexible File Format for Solid Freeform Fabrication," in: Solid Freeform Fabrication Symposium Proceedings, L. Marcus, 1. J. Beaman, J. W. Barlow, D. L. Bourell, and R. H.

1. Rock and M. J. Wozny, "A Flexible File Format for Solid Freeform Fabrication," in: Solid Freeform Fabrication Symposium Proceedings, L. Marcus, 1. J. Beaman, J. W. Barlow, D. L. Bourell, and R. H. Crawford (eds.), The University of Texas at Austin (Austin, TX), August 12-14 1991, pp. 1-12.[3]1. Arline, "A user survey of problems with rapid prototyping systems," Prototyping Report, vol. 1, no. 6, November 1991, pp. 7-8.[4]S. J. Rock, "Solid Freeform Fabrication and CAD System Interfacing," Master's thesis, Rensselaer Polytechnic Institute, Troy, NY, December 1991.[5]R. A. Schubert, "StereoLithography," in: NCGA'89 Conference Proceedings, vol. 2, (Philadelphia, PA), April 17-20, 1989, pp. 182-186.[6]C. R. Deckard, Selective Laser Sintering (CAD/CAM). Ph.D. thesis, The University of Texas at Austin, 1989.[7]B. Rooney. Private communication, Brock Rooney & Associates, Birmingham, MI, May 27, 1992.[8]C. Alexander.

Ph.D. thesis, The University of Texas at Austin, 1989.[7]B. Rooney. Private communication, Brock Rooney & Associates, Birmingham, MI, May 27, 1992.[8]C. Alexander. Private communication, 3D Systems, Inc., Valencia, CA, May 29, 1992.[9]S. J. Rock and M. J. Wozny, "Utilizing Topological Information to Increase Scan Vector Generation Efficiency," in: Solid Freeform Fabrication Symposium Proceedings, H. L. Marcus, 1. 1. Beaman, 1. W. Barlow, D. L. Bourell, and R. Crawford (eds.), The University of Texas at Austin (Austin, TX), August 12-14 1991, pp. 28-36.[10] M. MantyHi, "Boolean Operations of 2-Manifolds through Vertex Neighborhood Classification," ACM Transactions on Graphics, vol. 5, no. 1, January 1986, pp. 1 29.[11] K.J. Weiler, Topological Structures for Geometric Modeling. Ph.D. Rensselaer Polytechnic Institute, Troy, NY, August 1986. (RDRC-TR 86032).[12] F. S. Roberts, Applied Combinatorics. Prentice-Hall, 1984.94thesis,.

tomatic Feedback Control in Electron Beam Melting Using InfraredThermographyJorge Mireles1, Cesar Terrazas1, Francisco Medina1, Ryan Wicker1 1. W.M. Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, TXAbstractAn infrared (IR) camera has been installed in an Arcam A2 Electron Beam Melting (EBM) system for improved layer-by-layer monitoring and feedback control of the EBM build process. Previous work has demonstrated the temperature variations possible during a build (e.g., part/powder bed temperature elevates as build height increases) that have been shown to produce microstructural changes as well as a range of defects that can be detected (e.g., temperature anomalies and porosity). A stabilization of temperature during a build can lead to more uniform microstructure and mechanical properties throughout the fabricated part.

A stabilization of temperature during a build can lead to more uniform microstructure and mechanical properties throughout the fabricated part. Further, full spatial and temporal control of temperature could lead to controlled microstructural architectures in EBM-fabricated parts. An automatic feedback control system was developed to acquire a temperature matrix of the current layer and used as an input to a ‘ghost operator’ that modifies the necessary parameters (speed function, beam current, melt cycle’s post-heating time) for temperature stabilization to minimize microstructural variations.IntroductionThe Electron Beam Melting (EBM) process has been a promising technology for the rapid manufacturing of end-use metal components. Several materials have been processed using EBM (e.g. Ti-6Al-4V, Inconel 625, and Inconel 718) that are of interest in the aerospace industry for flight-ready part fabrication (Rodriguez et al., 2012, Murr et al., 2011).

Ti-6Al-4V, Inconel 625, and Inconel 718) that are of interest in the aerospace industry for flight-ready part fabrication (Rodriguez et al., 2012, Murr et al., 2011). Although part fabrication using Ti-6Al-4V has resulted in dense parts with mechanical properties comparable in microstructure still exist, namely differences in microstructure throughout a part that are due to a thermal gradient within the build chamber during fabrication (Murr et al., 2009; Puebla et al., 2012). Such uncontrolled variations are not suitable when mechanically isotropic parts are needed.

Such uncontrolled variations are not suitable when mechanically isotropic parts are needed. Furthermore, if a higher level of control is present in EBM, it may be possible to fabricate mechanically isotropic parts or parts with prescribed anisotropy, ensure manufacturing repeatability and reproducibility, and meet an adequate level of quality.to wroughttitanium, several anomaliesA FLIR SC645 infrared (IR) camera was installed in an Arcam A2 EBM system for the purpose of layer-by-layer monitoring and feedback (Rodriguez et al., 2012). Studies were performed observing parts produced using Ti-6Al-4V and results included the detection of defects such as porosity, non-uniform temperature between parts in a single build, non-uniform temperature within a single part, and over-melt of powder. Using the improved feedback from the IR system, an operator was able to manually modify parameters to achieve a more uniform build temperature (Rodriguez, 2012).

Using the improved feedback from the IR system, an operator was able to manually modify parameters to achieve a more uniform build temperature (Rodriguez, 2012). The current research focuses on developing an automatic feedback control method in an attempt to achieve a uniform build temperature to attain microstructural uniformity, detect defects, and achieve a layer-by-layer control of parameters to change grain size in Ti-6Al-4V microstructure.ConceptMicrostructural gradients in EBM-fabricated parts have been previously studied. These results show grain size differences from the bottom of a build to the top of a build where the bottom of a build exhibits a finer microstructure than the top of a build (Murr et al., 2009). Differences in grain size have been attributed to changes in mechanical properties such as hardness, ultimate tensile strength, and elongation (Murr et al., 2009, Puebla et al., 2012).

Differences in grain size have been attributed to changes in mechanical properties such as hardness, ultimate tensile strength, and elongation (Murr et al., 2009, Puebla et al., 2012). Moreover, some microstructural differences still exist after performing a HIP operation (Mireles et al., 2013). The layer-by-layer control achieved using an automatic feedback control loop can either attempt to stabilize process temperature to achieve uniform grain size or attempt to purposely increase/decrease process temperature through parameter modifications in an attempt to achieve a desired grain size or grain size gradient part. The basic concept for grain growth is described by the following grain growth equation (Porter & Easterling, 1981):𝟐 = 𝑫𝟎𝟐 + 𝑲𝒕𝑫𝟐where D2 is the final grain size, D0 is the initial grain size, K is a constant relating heating temperature and activation energy for grain growth, and t is time.

Although these fundamental equations are formulated to apply to single-phase metals, studies on Ti-6Al- 4V by Gil and Planell, 2000, showed such kinetics for grain growth apply for this alloy. As a result, as temperature rises, growth rate increases, and further increases with time when subjected to a high temperature (Gil & Planell, 2000). Additionally, performing heat treatment operations such as HIPing of EBM fabricated Ti-6Al-4V parts has shown to increase grain size due to the exposure of a high-temperature cycle that allows grain growth (Al-Bermani et al., 2010). Thus, grain growth is both time and temperature dependent.MethodsImage acquisition after the melt of each layer using an IR camera was previously achieved through the installation of a shutter mechanism and the use of National Instruments hardware and virtual instruments from LabVIEW software (Rodriguez, 2013).

In the present work, a virtual instrument was developed to perform automatic control from feedback obtained by images from the IR camera and layer information deciphered from calculations made by the virtual instrument’s loop iterations. The lack of a direct application programming interface (API) restricted direct communication with the Arcam EBM Control software to interact and change parameters programmatically on-demand. Therefore, applications were developed which triggered mouse click events to interact with the EBM Control interface using Microsoft Visual Basic 2010, consisting of (1) a program that allowed mapping the X and Y coordinates of different elements within the user interface (buttons, menus, lists, etc.) and (2) a routine that read the mapped coordinate values and triggered the mouse clicks. When called by the image acquisition virtual instrument, the simulation of mouse clicks began to change each parameter accordingly.

When called by the image acquisition virtual instrument, the simulation of mouse clicks began to change each parameter accordingly. Two computers were used due to the large memory consumption of image processing that may cause EBM system failure. Communication between the two computing systems was achieved through an Ethernet protocol.The developed application relies on data from a log file that must be initialized by the operator prior to the start of an experiment or build. Missing this step will cause anexception error when calling the mouse-triggering application potentially crashing execution of the entire monitoring system. An even simpler source of error can result if the computer mouse is manually moved while the application is called. This will give inaccurate coordinate values to the application resulting in unchanged or falsely changed parameters. Furthermore, the application will not have any effect if the computer screen is locked.

Furthermore, the application will not have any effect if the computer screen is locked. Finally, the mapping of the coordinates is specific to the monitor of the Arcam A2 machine used. Therefore, recalibration will be necessary if this application is to be used in another EBM system.Figure 1 Virtual Instrument for image acquisition and processingThe image acquisition virtual instrument is shown in Figure 1 and requires basic inputs from the operator including number of pictures per layer desired and its corresponding time delay from the instance the melt step has finished (images are taken after a melt step has ended). Included in the virtual instrument were options to run a grain size control algorithm, a temperature stabilization algorithm that performed image processing to detect parts and read average temperature of parts, and a porosity detection algorithm that used previous acquired images of porosity in EBM-fabricated parts as a reference for detection.

Six operator inputs were required to run temperature stabilization including upper and lower limits to account for IR measurement error (typically a desired temperature ± 30-40°C), a layer height to begin stabilization (typically ~2mm to account for initial temperature decay), a layer buffer (typically 6-10 to allow for parameters to take effect), and amount of post-heating time increase/decrease for parameter changes when temperature was found to be outside the desired limits.Using the image acquisition virtual instrument for grain size control, a 16mm diameter by 30mm long cylindrical specimen was fabricated as shown by Figure 2(a). The virtual instrument utilized a counter within its loop structure to compute the current layer and compared it against the user input for desired height for changes to take effect.

The virtual instrument utilized a counter within its loop structure to compute the current layer and compared it against the user input for desired height for changes to take effect. To obtain a part containing a grain size gradient, a single parameter for the melt cycle post-heating time was increased at 10mm and 20mm from the standard of 25seconds to 45seconds and again from 45seconds to 15seconds, respectively.Figure 2 Fabricated specimens and cut sections for a) grain size control and b) temperature stabilization where the red faces are the sections whose microstructure was analyzedA second cylinder (16mm diameter by 20mm long) was fabricated using a different build to test the temperature stability algorithm and porosity detection algorithm (Figure 2(b)). To perform temperature stability and porosity detection, an image processing step was utilized that performed an image conversion step from the initially grayscale image into a binary image.

To perform temperature stability and porosity detection, an image processing step was utilized that performed an image conversion step from the initially grayscale image into a binary image. A part was detected from a pre-defined region of interest and segmentation was achieved by applying an intensity threshold that was defined using previously acquired images of dark intensity level of 195. For temperature stabilization, an average temperature was recorded for the detected part that was used for comparison against the user’s temperature range input. The porosity algorithm also used segmentation procedures by applying an intensity threshold to a bright intensity level of 185 that has also been previously defined by images containing porous parts. A numerical amount of porosity was detected and porosity area was measured based on a calibration performed in the algorithm that took pixel size (0.175mm) and computed a total area.

A numerical amount of porosity was detected and porosity area was measured based on a calibration performed in the algorithm that took pixel size (0.175mm) and computed a total area. The amount of porosity that was measured can be used to compare against a user defined “acceptable range of porosity” in parts (e.g. stop a build if porosity has been detected or continue fabrication if a small amount of porosity is acceptable).After fabrication, the cylinders were cut using an IsoMet 400 Precision Saw (Buehler, Lake Bluff, IL) into three sections (Figure 2) including a section from the bottom corresponding to a standard post-heating time of 25 seconds for the melt step, a middle section with an elevated post-heating time of 45 seconds, and a top section with a lower post-heating time of 15 seconds. Cylinders fabricated to test the temperature stabilization algorithm were analyzed at the top and bottom sections ~10mm apart as shown by Figure 2.

Cylinders fabricated to test the temperature stabilization algorithm were analyzed at the top and bottom sections ~10mm apart as shown by Figure 2. After cutting, the sectioned pieces were mounted with a KoldMount(CMP Industries, Inc., Albany, NY) specimen mounting resin. Metallography was performed using 80grit, 320grit, 500grit, 800grit, 1000grit, and 1200grit, followed by a finishing step with a polishing cloth using 0.1µ and 0.05µ alumina slurry. An etchant composed of 100ml distilled water (H2O), 5ml nitric acid (HNO3), and 2.5ml hydrofluoric acid (HF) was applied for ~15seconds to reveal the microstructure.

An etchant composed of 100ml distilled water (H2O), 5ml nitric acid (HNO3), and 2.5ml hydrofluoric acid (HF) was applied for ~15seconds to reveal the microstructure. Optical microscopy was performed on the mounted pieces using a Leica MEF4M optical microscope with a digital imaging system.ResultsTemperature measurements were taken from IR images and recorded versus time as shown in Figure 3(a) where the sections are labeled corresponding to the standard build’s post-heating time parameter, grain coarsening section obtained during the increased post-heating time parameter, and grain refinement section corresponding to the decreased post-heating time parameter. IR images for each section were taken and are shown in Figure 3(b) displaying differences in IR temperature obtained by changing heat time parameters. Each measurement is an average of a cylinder’s melt area and the variations shown by Figure 3(a) can be due to porosity that may be inside the measurement area.

Each measurement is an average of a cylinder’s melt area and the variations shown by Figure 3(a) can be due to porosity that may be inside the measurement area. Figure 4 shows the results for the microstructural analysis performed showing standard grain structure when a) standard post-heating time parameters were utilized, coarsened grains when b) increased post-heating time parameters were utilized, and refined grains at c) where lowered post-heating time parameters were utilized. All optical images in this study correspond to acicular α-plate Widmanstätten-like microstructure (Murr et al., 2009). Measurements pertaining to the α grain width for the standard post-heating time parameters were ~0.65µm, ~1.31µm for the elevated post- heating time, and ~0.96µm for the lowered post-heating time.

Measurements pertaining to the α grain width for the standard post-heating time parameters were ~0.65µm, ~1.31µm for the elevated post- heating time, and ~0.96µm for the lowered post-heating time. Although the post-heating time setting of the top section was reduced below that of the standard, a smaller grain size was not achieved due to the thermal gradient that caused α grain growth in vertically fabricated parts (Puebla et al., 2012). A further reduction in grain size may not have been achieved since other factors such as build time or part size may affect grain growth. Volume of heat for a build can vary from build to build and layer-to-layer thus grain growth for varying build heights, overall part size, and/or heating/cooling time may contribute to grain size variations.

Nonetheless, it is noteworthy that the final grain size remained below ~1.31µm and further coarsening was not evident.Figure 5 shows the functioning temperature stabilization interface displaying part detection and temperature measurements. A conditional statement was used in the algorithm to change parameters when the temperature was detected beyond the specified limits. It continued building normally when the temperature was detected within the specified temperature limits. Figure 6 shows the microstructure achieved after running the algorithm. Measurements pertaining to the α grain width are 0.62µm for the bottom and 0.98µm for the top.

Figure 6 shows the microstructure achieved after running the algorithm. Measurements pertaining to the α grain width are 0.62µm for the bottom and 0.98µm for the top. Under the rationale that a stable build temperature should result in a constant grain size, the parameter modifications performed were not sufficient to stabilize the build temperature as evidenced from a ~0.36µm difference in grain size fromFigure 3 Grain growth experiment showing a) temperature graph from IR images and b) IR image comparison for each standard, coarsened, and refined grain sections where red spots on melt area are porosity that occurred during fabricationFigure 4 Microstructure images of grain sizes for each section a) standard post-heating time, b) elevated post-heating time parameters, and c) lowered post-heating time parametersthe bottom to the top of the part.

Figure 7 shows the temperature measurements taken from the IR images for the temperature stabilization experiment and still shows a slight climb in temperature through fabrication. Several factors may be attributed to the inability to acquire grain uniformity such as uncertainty in the IR temperature measurements that gives variable temperature readings from one layer to the next as well as build temperature variations that may be present in IR images resulting from the build process (Rodriguez, 2013). Furthermore, the changes were only performed by increment/decrement of 2seconds thus it may not be sufficient to produce the desired temperature change. In addition, although the part continued building for both the temperature stabilization and grain growth algorithm, parts exhibited porosity after several parameter changes that may require either a different parameter or a group of parameters to be changed.

A method that can help achieve uniform microstructure from the bottom of the build to the top of the build is to achieve a higher level of build chamber temperature control where either temperature is lowered as a build progresses or the build chamber temperature is maintained constant throughout a build.Figure 5 Functioning temperature stabilization interface showing part detection and temperature measurementsIntensity levels for previously identified porosity in IR images for EBM- fabricated parts has been measured and associated to object intensity levels in the virtual instrument’s object detection algorithm to identify porosity during fabrication. Porosity in IR images can be seen in the fabricated parts shown in Figure 3 (b) for the standard and coarsened sections and are described by the red spots inside the melt area. Using the intensity threshold technique, porosity can be detected as shown in Figure 4.

Using the intensity threshold technique, porosity can be detected as shown in Figure 4. Based on the porosity measurements, a build can be stopped after potential porosity is detected or after a critical area of porosity is reached. Porosity detection is documented throughout the build and given to the operator for analysis of instances where porosity was detected that is useful for post-build analysis.Figure 6 Microstructure of temperature stabilization algorithm for a) bottom section and b) top sectionFigure 7 IR temperature results for temperature stabilization experimentConclusionsUsing a previously installed IR camera in an Arcam A2 EBM system an improved image acquisition virtual instrument was created using LabVIEW software (Rodriguez, 2013).

The improved virtual instrument allowed for automatic control of EBM technology to 1) achieve parameter modification using simulated mouse clicks useful for grain size control, 2) attempt temperature stabilization by recording a part’s average temperature using image processing and automatically making decisions, 3) detect porosity during a build to determine if part fabrication needs to be stopped or porosity detection to be used for post-build analysis. Results show a grain size gradient wasachieved within a part but other part properties were also affected such as the development of porosity. Part detection and average temperature readings were taken to achieve temperature stabilization; however, development of part porosity presented a part quality problem when changing a single build parameter.

Porosity in fabricated parts was successfully detected and can be used as a notification tool to a user in case an unacceptable level of porosity is reached and fabrication needs to stop. Porosity information can be used in the future to correct porosity via the control algorithm by parameter changes.Future WorkAlthough a grain size gradient was achieved, it is still required to document parameter changes that produce a specific gain size. As previously mentioned, both time and temperature variables affect final grain size, therefore, variables affecting time that a part is left in the machine (e.g. height of build, heating/cooling time, overall size of part) may affect overall grain size and needs to be considered. For all automatic control operations, modification of parameters through simulated mouse clicks is not ideal since changes do not take effect immediately.

For all automatic control operations, modification of parameters through simulated mouse clicks is not ideal since changes do not take effect immediately. Direct access to Arcam software would provide for instantaneous parameter changes instead of using mouse click simulations. Two computers were used in this study due to the high memory consumption of image processing that can affect computer processing causing build failure. Therefore a more powerful computer would be needed to mitigate possible communication issues or any communication delays that may occur.References[1] Rodriguez, E., (2013). Development of a thermal imaging feedback control system in electron beam melting. Master’s thesis for master’s degree, University of Texas at El Paso, El Paso, Texas, USA.[2] Schwerdtfeger, J., Singer, R.F., Korner, C., (2012). In situ flaw detection by IR- imaging during electron beam melting. Rapid Prototyping Journal, 18 (4), pp.

tomatic Generation of Strong, Light, Mutli-Functional Structures from FEA Outputand D. Stahl†D. Cook, B. Knier, V. GervasiApplied-Technology Center, Rapid-Prototyping Research, Milwaukee School of Engineering, Milwaukee, WI 53202 †Department of Architectural Engineering and Building Construction, Milwaukee School of Engineering, Milwaukee, WI 53202 Abstract An automated process is being developed that will generate a minimal-mass lattice structure that is fabrication-ready for a selected solid-freeform-fabrication (SFF) process. The results of a standard, structural, finite-element analysis (FEA) are processed to define the selection, alignment and sizing of unit lattice elements, such that a minimal-mass structure can be defined. This process will allow for considerations of structural performance (e.g. safety factor), multiple loads, as well as process parameters (e.g. materials and min./max. sizes).

This process will allow for considerations of structural performance (e.g. safety factor), multiple loads, as well as process parameters (e.g. materials and min./max. sizes). Further development would lead to the definitions of composites and multi-functionality, as well as high-performance- computing (HPC) capabilities.I.IntroductionA. PurposeThe purpose of this research was to determine an effective means of defining fabricatable geometry, representative of those results that may be obtained through structural optimization, for the design and fabrication of compact, efficient and effective fluid-power components.While this research was conducted specifically for the advancement of fluid-powertechnologies (The Center for Compact and Efficient Fluid Power), the algorithms developed and knowledge gained may be applied to nearly all industries.

In particular, the biomedical and aerospace industries also have immediate need for such variable-density, minimal-mass, multi- functional component designs. Apprehension to the adoption of additive-manufacturing processes for these industries necessitates the development of standards, such as ASTM’s Committee F42 on Additive Manufacturing Technologies (ASTM International, 2010). One goal of this paper, therefore, is to convey the potential of these technologies to designers and manufacturers so that they may be further motivated to bring these capabilities to the mainstream.A second goal is to further contribute to the growing community that seeks to capitalize onthe combination of capabilities afforded by high-performance computing (HPC), multi- disciplinary design optimization (MDO) and additive manufacturing (AM).B. Background1.

Background1. Design Optimization:Design optimization is a process of considering design alternatives having properties within the specified constraints, and selecting the design that most-closely matches the objective, within the allowable tolerance. For component designs, this can be achieved through structural optimization.Topology optimization is the most common structural-optimization algorithm for defining complex three-dimensional designs.

The component is contained within a design volume (far left) and boundary conditions are applied (mid left). The output is a gradient-density definition of the optimal design (mid and far right).This optimization approach is also limited in the resolution of its computed optimum. The minimal feature cross-sectional dimension must be larger than the minimum mesh size, typically by a factor of five. Altair’s Optistruct® algorithm also imposes a maximum ratio between maximum and minimum element sizes of 5:1. Therefore, for multi-functional geometries requiring high resolution, i.e. a high surface-area-to-volume ratio, the state-of-the-art requires extremely fine meshes throughout the entire design domain, or locally via iterative mesh refinements.

a high surface-area-to-volume ratio, the state-of-the-art requires extremely fine meshes throughout the entire design domain, or locally via iterative mesh refinements. Both of these discretization approaches are computationally expensive over the course of the optimization.The Homogenization approach is another structural-optimization algorithm, and explicitly considers the generally non-isotropic optimal structure; however, this approach has not yet been realized, to the authors’ knowledge, for three-dimensional designs, and generally only considers orthogonal structures, limiting the optimization potential (Bendsoe & Kikuchi, 1988) (Hassani & Hinton, 1999). Shear stresses, arising prevalently in multiple-load-case optimizations, necessitate the oversizing of the beams, struts and joints in orthogonal structures.1 FE TOSCA Structure (FE-DESIGN GmbH, Technologiefabrik Bau A, EG.) is another commercial software; but, it’s algorithm has not been confirmed as SIMP.2.

Multi-Functionality of Structures:In addition to bearing mechanical load using minimal mass, engineered structures can also perform other functions, allowing for the design of multi-functional components and systems. Of particular interest, for the fluid-power industry and others, is the potential for metal-based structures to provide thermal management and even acoustic-pressure absorption. While these variable-density structures may be quite complex, their fabrication is now possible with additive- manufacturing processes.3. Structure-Generation Options:Once the optimized geometry is defined, the structure of corresponding properties must be generated within. The known algorithms for automatically generating lattice structures within geometric constraints have been grouped as follows:a.

The known algorithms for automatically generating lattice structures within geometric constraints have been grouped as follows:a. Space-filling latticeThese algorithms simply copy a base unit-lattice geometry in their respective coordinate systems, typically in the Cartesian system, to the defined extents limits or specified number of units. Modifications to the geometry among the units must be performed manually. Fitting these structures within an irregular bounding geometry requires a Boolean intersection operation. The options are:(cid:1) Selective Space Structures® (Netfabb GmbH, 2010) (cid:1) TetraShell™ (Editors, DE, 2008) (cid:1) Conventional computer-aided-design (CAD) software (cid:1) Conventional modeling softwareb. Conformal latticeThese algorithms place the structural junctions along the surface of theprescribing/bounding geometry, thereby conforming to its curvature, etc. Typically, the nodes of a surface mesh define these coordinates.

Typically, the nodes of a surface mesh define these coordinates. Successive layers subdivide the space between the outer bounding surfaces, and have corresponding junction points for simple, point-to-point connectivity of the structure. Because each unit is dynamically defined, the geometry of each may be automatically modified by the selected algorithm. The options are:(cid:1) Truss Creator® - Georgia Tech. Systems' Realization Lab® – (Wang, 2005) (cid:1) Paracloud® (generative modeling) – requires pre-defined mapping (cid:1) Scripting within modeling software, e.g. Python® scripting within Blender®II. ApproachA. “Stress-Field-Compliant” ApproachStahl’s approach was to physically realize the model used in the homogenization approach through AM processes (Stahl & Batdorff, 2004). Additionally, for maximum optimization, the resultant data was processed such that a “stress-field-compliant” mesh could be defined.

Additionally, for maximum optimization, the resultant data was processed such that a “stress-field-compliant” mesh could be defined. This mesh connectivity, then, served as the connectivity for the mutually-orthogonal lattice structure that was to be generated (Figure 2).Figure 2: Stress-field-compliant homogenization method (mid left) principal stresses (middle). The orthogonal “ quadrilateral element, thus aligned with the stress vectors; and the struts were sized ; and the struts were sized quadrilateral element, thus aligned five layers in the third planar structure was then stacked five layers in the third accordingly. The resultant planar structure was (Stahl & Batdorff, 2004). dimension (right) (Stahl & Batdorff, 2004)lattice generation. The density distribution of the compliant-lattice generation. The density distribution of the to align the elements with the (mid left) is re-meshed (mid right) to align the elements with the inserted into each.

The density distribution of the to align the elements with the (mid left) is re-meshed (mid right) to align the elements with the inserted into each . The orthogonal “Jack” unit lattice was inserted into eachB. “Onioin-Skin ApproachSeeking to expand Stahl’s work(Gervasi & Stahl, 2004), (Stahl & ), and structurally , a cube was loaded under shear, at diagonal corners (Figure 3), and structurally . A mid-rangeStahl’s work into the third dimension (Gervasi & Stahl, 2004)Batdorff, 2004), a cube was loaded under shear, at diagonal corners (Fig optimized using Altair’s Hyperworks® suite optimized using Altair’s Hyperworks® suite (Altair Engineering, Inc., 2007).

A mid iteration, one having a wide gradient of densities, was selected; and, this gradient was subdivided iteration, one having a wide gradient of densities, was selected; and, this gradient was subdivided iteration, one having a wide gradient of densities, was selected; and, this gradient was subdivided density, every ten percent, between 30% and 100%. For three-dimensional density, every ten percent, between 30% and 100%. For three into levels of percent-density, every ten percent, between 30% and 100%. For three objects, these levels are similar toevels are similar to the layers of an onion, hence the name.These individual layers were then exported as stereolithography (STL) files for fabrication. These individual layers were then exported as stereolithography (STL) files for fabrication. These individual layers were then exported as stereolithography (STL) files for fabrication.

These individual layers were then exported as stereolithography (STL) files for fabrication. These individual layers were then exported as stereolithography (STL) files for fabrication. Engineering, Inc., meshing these layers, within Altair’s Hypermesh® (Altair Engineering, Inc., Unfortunately, re-meshing these layers, within Altair’s Hypermesh® 2007), to align the elements with the principal stress directions, proved to be an insurmountable , to align the elements with the principal stress directions, proved to be an insurmountable , to align the elements with the principal stress directions, proved to be an insurmountable Magics RP® (Materialise, 2010) was used to realize these geometries as realize these geometries as task. As an alternative, Magics RP® Boolean intersected with a block density lattice structures. Each layer’s geometry was Boolean intersected with a block variable-density lattice structures.

As an alternative, Magics RP® Boolean intersected with a block density lattice structures. Each layer’s geometry was Boolean intersected with a block variable-density lattice structures. of lattice geometry, having the desired density to match the layer, leaving a section of lattice of lattice geometry, having the desired density to match the layer, leaving a section of lattice of lattice geometry, having the desired density to match the layer, leaving a section of lattice matching the bounding geometry of the layer. The Tetralattice® struct (Gervasi V. R., 2001) matching the bounding geometry of the layer. The Tetralattice® structure (Gervasi V. R., 2001) was used here. Aligning the individual lattice layers and conducting a Boolean union, the was used here. Aligning the individual lattice layers and conducting a Boolean union, the was used here.

Aligning the individual lattice layers and conducting a Boolean union, the was used here. Aligning the individual lattice layers and conducting a Boolean union, the was used here. Aligning the individual lattice layers and conducting a Boolean union, the representing the results of the wise,” gradient-density geometry, representing the results of the resultant geometry was a “step-wise,” structural optimization (Figure 3) The component was fabricated using the Selective Laser ). The component was fabricated using the Selective Laser Sintering™ (SLS™) machine in the Rapid Prototyping Center™ (RPC™) of the Milwaukee Sintering™ (SLS™) machine in the Rapid Prototyping Center™ (RPC™) of the Milwaukee Sintering™ (SLS™) machine in the Rapid Prototyping Center™ (RPC™) of the Milwaukee School of Engineering (MSOE).Figure 3: "Onion-Skin" approach to generating variable from the gradient-density output of conventional topology optimization.

g density gradient is constrained cube (left) is structurally optimized; and, the resulting density gradient is constrained cube (left) is structurally optimized; and, the resultin . These layers are cut discretized into density ranges, like the layers of an onion (middle). These layers are cut discretized into density ranges, like the layers of an onionmass structures Skin" approach to generating variable-density, reduced-mass structures The loaded and density output of conventional topology optimization.

The loaded andand joined into one structure out from blocks of TetraLattice® of corresponding density, and joined into one structure out from blocks of TetraLattice® of corresponding density (right).While the “Onion-Skin” process demonstratethe feasibility of directly defining a three- Skin” process demonstrates the feasibility of directly defining a three mass, solid-freeform-fabricatable, variable-density structuredensity structure: the process is dimensional, reduced-mass, solid the resultant geometry is not completely optimized, i.e. the lattice elements are , i.e. the lattice elements are not automated; the resultant geometry is not completely optimized is weakened with stress concentrators at the directed or sized; and, the structure is weakened with stress concentrators at not stress-field-directed or sized; sectional areas. A conformal layer interfaces, due to the instantaneous changes in cross-sectional areas.

A conformal layer interfaces, due to the instantaneous changes in cross-sectional areas. A conformal layer interfaces, due to the instantaneous changes in cross s concentrations by structure, one having junctions along these interfaces, could avert these stress concentrations by structure, one having junctions along these interfaces, could avert these stres designing the proper geometry at each junction. designing the proper geometry at each junction.C. “Stress-Field-Directed” Approach ” Approachdominate load cases, the Von Mises stress The observation was made that, for such shear-dominate load cases, the Von Mises stress The observation was made that, for such shear field closely approximated the end results of the topology optimization process. In an attempt, optimization process.

In an attempt, optimization process. In an attempt, field closely approximated the end results of the topology the often high computational costs, the Von Mises therefore, to shortcut the process, avoiding the often high computational costs, the Von Mises therefore, to shortcut the process field, from a single finite-element analysis lie at regular intervals along the iso basic “Cube” lattice unit cell (Figure lattice unit cell (Figure 6) to make mutually-orthogonal connections lattice elements based on the enclosed area, a spider-web-like structure (Figure 4 lattice elements based on the enclosed area, a s for this classic Michelle load case (Bendsoe & Sigmund, 2004). This semi generated for this classic Michelle load case process: reads in the FEA output reads in the FEA output; determines the connectivity; sizes the lattice elements then generates the structure within modeling software, e.g. Rhinoceros®, via a script.

Rhinoceros®, via a script. then generates the structure withi while reduced in mass, this structurethe lattice junctions would element analysis, was processed; such that, the lattice junctions would). Using the s along the iso-surfaces of the stress field (Figure 4, middle). Using theorthogonal connections, and sizing the4, right) was . This semi-automatedsizes the lattice elements; and, Rhinoceros®, via a script. However,this structure is still not completely optimized.Figure 4: Semi-automatic generation of a stress (right) from the FEA output (middle)lattice structure generation of a stress-field-directed, and sized, lattice structure(left). (middle) of the classic Michelle cantilever (left)D. “Non-Iterative Optimization” Approach Iterative Optimization” ApproachThe realization was then made that a single FEA analysis (per load case) provides all of the.

“Non-Iterative Optimization” Approach Iterative Optimization” ApproachThe realization was then made that a single FEA analysis (per load case) provides all of the . This data could be , including specification of to estimate the distribution of stress requiring the least mass, including specification ofmade that a single FEA analysis (per load case) provide to non-iteratively define the optimal structure. This data couldnecessary load information to non processed to estimate the distribution of stress the lattice spacing. Knowing this definitions of the required unit lattices could be made.

This data couldnecessary load information to non processed to estimate the distribution of stress the lattice spacing. Knowing this definitions of the required unit lattices could be made. definitions of the required unit lattices could be made.stress distribution, including total and principal stress, local Knowing this stress distribution, including total and principal stress, localIn contrast to conventional optimization routines, this is a more “direct” approach, and, In contrast to conventional optimization routines, this is a more “direct” approach, and, In contrast to conventional optimization routines, this is a more “direct” approach, and, ues (once the initial FEA has successfully converged). therefore, does not suffer convergence issues (once the initial FEA has successfully converged).

therefore, does not suffer convergence issues (once the initial FEA has successfully converged). therefore, does not suffer convergence issfor a multi- This process cannot be disclosed here; but, a conceptual model of the process for a multi This process cannot be disclosed here; but, a conceptual model of the process functional, i.e. load-bearing and heat using the “Cube” unit-cell, the resulting structure looks like an inter webs. The suspicion is that this is not merely coincidental. webs. The suspicion is that this is not merely coincidental.bearing and heat-dissipating, tensile rod is shown in Figure 55.

The suspicion is that this is not merely coincidental. webs. The suspicion is that this is not merely coincidental.bearing and heat-dissipating, tensile rod is shown in Figure 55. Once again, connected stack of spider cell, the resulting structure looks like an inter-connected stack of spiderFor further mass reduction, automated selection from the available unit lattice types is For further mass reduction, automated selection from the available unit lattice For further mass reduction, automated selection from the available unit lattice dependent functional definition of this structure also requires geometry-dependentnecessary. Multi-functional definition of this structure also requires geometry characterizations of these unit lattices for those functions.

Multi-functional definition of this structure also requires geometry characterizations of these unit lattices for those functions. characterizations of these unit lattices for those functions.Figure 5: Point-loaded tensile rod of a load- and heat-dissipation-and the concept model of the automatic generation loaded tensile rod (left) and the concept model of the automatic generationoptimized structure (right).E. Structure Characterizationiterative generation of lattice structures, however, requires the constitutive This non-iterative generation of lattice structures, however, requires the constitutive iterative generation of lattice structures, however, requires the constitutive Hassani & Hinton showed structure- properties of all unit lattice structures to be employed. Hassani & Hinton showed structure properties of all unit lattice structures to be employed.

Hassani & Hinton showed structure properties of all unit lattice structures to be employed. uctures that were dimensional, orthogonal, variable-density structures that were performance plots for the two-dimensional, orthogonal, variable . These three-dimensional algorithms (Hassani & Hinton, 1999). These three used in their homogenization algorithms , i.e. elastic and shear the constitutive-matrix components, i.e. elastic and shear plots presented the variations in the constitutive ts, for one unit-lattice type. moduli, with the varying widths of the orthogonal structural elemen The algorithm under development dimensional unit lattices (Figurewidths of the orthogonal structural elements, for one unit multiple, three- under development, however, requires full characterizations of multiple, threein relation to their potential variations in geometry.

(Figure 4), in relation to their potential variations in geometry.considered in the automated structure- lattice unit cells considered in the automated structure Figure 6: Four of the five lattice Each hexahedral unit cell is comprised of eight mutually-orthogonal Each hexahedral unit cell is comprised of eight mutually generation algorithm.

From this strain and resultant-nodal-force data, for various geometric configurations of each unit lattice, their geometry-dependent constitutive relations could be determined.III. ResultsA. Unit-Lattice CharacterizationsPreliminary analyses indicated that the solid, Isotropic Material with Penalization (SIMP)approach, used by the two leading commercial software suites (Altair’s Optistruct® and VR&D’s Genesis®), did not accurately represent the performance of the simple orthogonal “Jack” unit lattice, or “Cube” unit cell (Knier, 2009). Shown in Figure 7, there are significant discrepancies between the actual unit-lattice properties and the exponential SIMP relation, necessitating the complete characterization of the unit lattices to be fabricated for accurate representation.Apparent Normalized Elasticity vs.

Volume Fraction for the "Jack" Unit Latticee s a b E / ) t n e r a p p a ( x E1 0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0CalculationSIMP, n=3SIMP, n=200.10.20.30.40.50.60.70.80.91Volume FractionFigure 7: Normalized Ex elasticity for the "Jack" unit lattice for a range of volume fractions resulting from varying combinations of orthogonal-strut diameters. Curves for the SIMP relation are also shown for comparison.Using nine different geometry configurations, i.e. x, y and z-strut diameters, the effectiveconstitutive components were tabulated within Microsoft Excel, that has an integrated tool,Linest, for multi-variate regression.

x, y and z-strut diameters, the effectiveconstitutive components were tabulated within Microsoft Excel, that has an integrated tool,Linest, for multi-variate regression. For example, assuming a 2nd-order, homogeneous, cross- multiplied polynomial for the “Super Cube" (Equation 3):222(3)E=2.12e7Φ xx+7.5e6Φy+7.47e6Φz+2.63e5ΦΦ xy+2.91e5ΦΦ x3.78Φ zΦ yzThis regression process was repeated for all of the orthotropic properties of each of the unit-lattice structures; E, E x, E y, G z, G, G, ν, νand ν. Table 1 shows a comparison of two ofyz the constitutive components, Ex and Gxy, for each of the five unit lattices under consideration for equal orthogonal geometries. The “Ultra” and “Super” cubes have the highest components for this selected configuration.xyyzxzxyzxTable 1: Comparative, effective, constitutive components for the five unit lattices of consideration having equal orthogonal strut diameters.

The Octet Truss has no vertical strut, hence zero.Type Cube ("Jack") Super Cube Ultra Cube Modified Super Cube OctetΦx (mm) Φy (mm) Φz (mm) Ex (GPa) Gxy (GPa) 33.1 20.32 65.7 20.32 65.5 20.32 36.3 20.32 32.0 05.08 5.08 5.08 5.08 5.0820.32 20.32 20.32 20.32 20.3255.3 168.3 178.1 95.8 85.5Note that these values must be considered along with the structure’s volume fraction. While the “Super Cube” is significantly stiffer than the “Cube,” it is also more massive. These are the considerations to be made by the automated unit-lattice selection and sizing algorithm that is currently under developmentIV. Next Steps1. Completion of AutomationThe automated selection, population and sizing of other unit lattice elements into conformal geometries will be finalized. Additionally, considerations for the limitations of the selected AM technology will be included, e.g. minimum feature dimensions, layer thickness, etc.2.

Additionally, considerations for the limitations of the selected AM technology will be included, e.g. minimum feature dimensions, layer thickness, etc.2. Continue CharacterizationsThe unit-lattice structures will continue to be characterized for their performance in heat conduction/convection and acoustic absorption. Additionally, two-material composites will be considered.3. Continue DevelopmentSupport will be sought to continue the development of the non-iterative optimization routines, to include; computational geometry, parallel processing, multi-disciplinary design optimization and meshless methods.4. Compare PerformanceComparative determination of performance gains for the new algorithm will then be made against topology optimization and genetic algorithms.5. Conduct TestingOnce components are designed and fabricated, physical testing is required to validate the design algorithm(s).V.

Conduct TestingOnce components are designed and fabricated, physical testing is required to validate the design algorithm(s).V. ConclusionsThe “Onion-skin” approach demonstrated the feasibility of defining variable-density geometry from topology optimization results; and, semi-automated structure definition and sizing were demonstrated with the “stress-field-directed” approach. Neither of these structures, however, was completely optimized.The non-iterative method holds the greatest potential for multi-disciplinary structuraloptimization; but, the algorithm must be completed before testing can be conducted. To this end, the unit lattices under consideration for this method must also be completely characterized for all anticipated functional requirements, such as; load bearing, thermal management and acoustic absorption.

Once this algorithm is completed, structures can be fabricated and physically tested; and, performance comparisons can be made against topology-optimization and genetic algorithms.VI. Aknowledgements(cid:2) This material is based upon work supported by the National Science Foundation’s Engineering Research Center, the Center for Compact and Efficient Fluid Power (CCEFP), under Grant No. EEC-0540834. Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author and do not necessarily reflect the views of the National Science Foundation.(cid:2) Altair Engineering (Chad Mach, Michael Dambach and R. Ramcharan) – Access toHyperworks® suite and technical support.(cid:2) Parker (Brewce Larkin) – Donation of pump, access to CAD models and modelingsupport.(cid:2) Research assistants (MSOE-ATC-RPR): Gunnar Vikberg, Aaron Kimball, Lee VanDehy, Adam Leslie, Andrew Steevens (REU ’07),(cid:2) Jay Urbain, Ph.D.

TOMATIC LAYERWISE ACQUISITION OF THERMAL AND GEOMETRIC DATA OF THE ELECTRON BEAM MELTING PROCESS USING INFRARED THERMOGRAPHY Ridwan, S., Mireles, J., Gaytan, S.M., Espalin, D., Wicker, R.B. W.M. Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, TexasAbstractLayerwise monitoring has become an area of interest in the field of additive manufacturing because of potential to further enable part qualification during every stage of fabrication. Spatial monitoring and qualification during part fabrication has never before been possible with traditional manufacturing processes such as milling or casting. An IR camera has been externally annexed atop an EBM system to obtain layerwise thermographs throughout the fabrication process. This paper demonstrates a process to compare each layer of fabrication using automatically acquired thermal images to the corresponding CAD file for each fabricated object.

This paper demonstrates a process to compare each layer of fabrication using automatically acquired thermal images to the corresponding CAD file for each fabricated object. Two different methods of image analysis for part detection were compared (analysis on the basis of color and analysis by edge detection). Detection allowed the quantification of processing information (average temperature and surface anomalies) and geometric information (surface area and perimeter). A percent error of the compared surface area was found to range from 5%- 17%, and automatically acquired temperature measurements were within 7.8K of the recorded thermograph.

A percent error of the compared surface area was found to range from 5%- 17%, and automatically acquired temperature measurements were within 7.8K of the recorded thermograph. The methods presented in this research showcase the beginning steps of integrated metrology in advanced manufacturing systems and automatic monitoring of per-part thermal behavior and part quality.1.0 IntroductionLayerwise monitoring has become an area of interest in the field of additive manufacturing (AM) because of potential to further enable part qualification during every stage of fabrication. Previous thermal monitoring of layerwise processes like electron beam melting (EBM) has allowed characterization of defects that ultimately affect microstructure and part geometry [1, 2]. Dinwiddie et al. demonstrated a method of continuous thermal imaging within the Arcam EBM system that allowed post-fabrication evaluation of prevalent defects, such as porosity. Yadroitsev et al.

Dinwiddie et al. demonstrated a method of continuous thermal imaging within the Arcam EBM system that allowed post-fabrication evaluation of prevalent defects, such as porosity. Yadroitsev et al. demonstrated an optical system for molten pool temperature measurements using a charge-coupled device camera for the selective laser melting process where brightness temperature distributions were analyzed and microstructure changes were observed for biomedical components. Craig et al. developed a multi-sensor thermal imaging suite to monitor a directed energy deposition process. The authors also developed a software feature to measure variations of thermal metrics. Although the above mentioned works allowed for improved AM system feedback, there is still a need to implement automatic monitoring during fabrication.In this research, a thermal camera has been installed atop an EBM system to monitor each layer after the melting process, or layer fabrication, has completed.

A method for detecting the size, shape, and surface area of a fabricated object was accomplished. Additionally, an algorithm was developed to compare the thermal data obtained from the installed camera with an input layer derived from CAD as a means to detect any prevalent porosity. Furthermore, the pixel values from the obtained thermographs were converted to temperature values and averaged to obtain and record an average temperature value within the object’s surface area. Such methodsof automatic temperature acquisition and layer-by-layer CAD to thermograph comparison showcase the beginning steps of integrated metrology in advanced manufacturing systems and automatic monitoring of per-part thermal behavior and part quality.2.0 MethodologyAM processes typically begin through a stereolithography (STL) file that is generated from a CAD model that works on the basis of triangulation of the three-dimensional geometry. STL files can be of two types: binary and ASCII.

STL files can be of two types: binary and ASCII. Precision of the geometry is affected by the layer thickness of the successive layers, or slices, typically ranged between 0.05-0.2mm [3]. The STL file is then converted to the format supported by the machine, or a slice file (two- dimensional file) of the STL geometry, that contains the beam path instructions. Although part geometry is currently used as input via a CAD model, various nebulous factors such as differences with the ideal surface area, shape, and the fabrication conditions of the manufacturing process (i.e. prevalence of defects such as warping) affect a part’s surface area.In this research, a FLIR SC645 infrared (IR) camera (FLIR Systems, Wilsonville, OR) was used to automatically output thermal images from each layer of a part’s surface area. The camera was tilted to 15 degrees with respect to the vertical axis of the fabrication chamber and views fabrication through a ZnSe glass.

The camera was tilted to 15 degrees with respect to the vertical axis of the fabrication chamber and views fabrication through a ZnSe glass. The installation of the IR camera in EBM is described in more detail in previous research work [3, 4]. A pneumatic system controls a shutter mechanism that protects the camera’s viewing window from metallization, or a coating phenomenon that occurs during melting [4]. A control system using LabVIEW (National Instruments, Austin, TX) was developed to control the shutter mechanism and automatically acquire thermal images of each layer after the melting process was completed [3, 4]. Voltage signals from the EBM system were acquired during the build process and interpreted by a modular controller (NI cRIO-9074, National Instruments, Austin, TX). These signals were used to automatically acquire thermographs from the build area.

These signals were used to automatically acquire thermographs from the build area. Using LabVIEW, thermographs were saved as bitmap images where each pixel value represented a temperature value after the melting process of the built object for each specific layer. Using the image and data acquired from the camera, a MATLAB (The MathWorks, Inc., Natick, MA) code was developed to obtain the desired image data, detect defects of an object, and find the average temperature value during/after the melting process.Using MATLAB, a region of interest was selected on the basis of the graphical input where four coordinates were selected manually at the beginning of image analysis. A matrix was formed on the basis of minimum and maximum values of X and Y coordinates to ensure that all fabricated objects were within the region of interest.

A matrix was formed on the basis of minimum and maximum values of X and Y coordinates to ensure that all fabricated objects were within the region of interest. The original image was then converted into a binary image, and boundaries were detected to identify the fabricated parts using two different techniques (edge detection and detection on the basis of color difference). The pixel values within the boundaries were read from the standardized image and converted to temperature values, and average temperature values were calculated for each fabricated part. Furthermore, the surface area (total number of pixels) and perimeter were obtained for each detected part.

Furthermore, the surface area (total number of pixels) and perimeter were obtained for each detected part. Finally, the ideal surface area for the fabricated object was measured from input layerwise data and a percent error from each surface area measured was calculated.StartDetect boundaries of the binary imageRead each pixel value within boundariesRead the imagePlot the imagePlot the boundaries with porosityConvert pixel values into temperature valuesTake graphical input to define region of interest (ROI)Remove small objectsPlot the boundaries without porosityCalculate average temperature value of each objectForm matrix for ROI dimensionsCrop ROIGet properties (area and perimeter) of the boundariesPrint the area and temperature of each detected boundaryConvert original image into binary imageFind minimum and maximum x and y coordinate of each object boundaryEndFigure 1.

Flowchart of steps for analysis of output layer based on color.2.1 AlgorithmsTwo different techniques were followed to analyze the output thermographs after the melting process: 1) analysis on the basis of color and 2) analysis on the basis of an edge detection technique. The steps in the analysis of the output layer on the basis of color are depicted in Figure 1.Analysis on the basis of color begins by taking the original image, cropping to segment fabrication area, and the detecting and labeling of boundaries for analysis. Pixels within the detected boundaries allowed the calculation of geometric variables such as surface area and perimeter as well as process variables such as average temperature values. For geometry, thearea of each pixel was previously calibrated [4] and used in this study for calculation. To calculate part temperature, pixel brightness was converted to a temperature value using the previously developed conversion method [3].

To calculate part temperature, pixel brightness was converted to a temperature value using the previously developed conversion method [3]. Upon calculation of necessary parameters, the code prints all the acquired data on-the-fly. Basic steps in edge detection technique utilized for this study are shown by Figure 2.StartConvert original image into grayscale imageRead the imageDetect edge of the grayscale imagePlot the imageJoin discontinuous edgesTake graphical input to get region of interest (ROI)Fill the boundariesForm matrix for ROI dimensionsPlot the boundaries within imageCrop ROIEndFigure 2. Flowchart of steps for analysis of output layer based on edge detection technique.The Sobel method, which is an image analysis method based on image intensity values, was used for edge detection. In this method, instead of detecting boundaries from the cropped color image obtained, the cropped image was converted to grayscale to facilitate edge detection.

In this method, instead of detecting boundaries from the cropped color image obtained, the cropped image was converted to grayscale to facilitate edge detection. Discontinuous edges were joined, and boundaries were filled to get an image of the fabricated object. Geometric and process variables were also capable of being obtained upon acquiring the image of the fabricated object.Although both of the above steps allow geometric measurement of fabricated parts, comparison of the ideal geometry is necessary for part qualification applications. In this study, measured geometry was compared to the ideal slice data from the input CAD file. A MATLAB code was developed to compare the input layer (ideal slice image) with the output layer (thermal image).

A MATLAB code was developed to compare the input layer (ideal slice image) with the output layer (thermal image). The steps for comparison are depicted in Figure 3.StartForm a file path name on the basis of the image number and formatCall for function that receives input image file path and gives surface areanCalculate the error percentage of areanNoPart fabrication is finishedCheck for file existenceYesCall for function that receives output image file path and gives surface areaOutput the error percentageEndEndFigure 3. Flowchart of steps for comparison of input and output layer.Comparison of the detected surface area starts by first matching the file name (corresponding to the layer being compared) corresponding to the current layer. If the file name for the ideal image matches that of the image being compared, comparison of surface area continues until fabrication is complete.

If the file name for the ideal image matches that of the image being compared, comparison of surface area continues until fabrication is complete. Here, two different script files were called to obtain the surface area for input and output layers. The first function script used the color difference technique (as described in Figure 1) and returned the surface area. In Figure 3, n is the number of iterations that depends on the number of layers of the built object as well as the number of images of the output layer after the object’s melting process has completed. Before the calculation of the error percentage, the size of the area matrix of both input and output layer images were also verified to ensure that the number of objects in both images were the same. A percent error was calculated upon comparison of both surface areas and recorded for each specific layer.3.0 Results and discussion:An example IR image (after cropping) is shown in Figure 4a.

A percent error was calculated upon comparison of both surface areas and recorded for each specific layer.3.0 Results and discussion:An example IR image (after cropping) is shown in Figure 4a. This represents the image used for detection on the basis of color (Figure 1). This original image is a representation of temperature, where lighter color represents higher temperature and darker color represents lower temperature (dependent on emissivity). Here, the build area of a particular layer, distinct (darker) low temperature regions from the thermal image can be seen within the solid cylinders being fabricated in this example. An example standardized grayscale image used for analysis on the basis of edge detection is shown in Figure 4b. The brightness differences shown by thea)b)Porosityc)d)Figure 4.

An example standardized grayscale image used for analysis on the basis of edge detection is shown in Figure 4b. The brightness differences shown by thea)b)Porosityc)d)Figure 4. Original image (a), standardized grayscale image (b), binary image of original image (c) and inverted binary image of standardized grayscale image (d).grayscale image (Figure 4b) represent temperature differences that, upon conversion, can be quantified as a temperature value. Figure 4c represents the binary image generated from the original thermal image, and Figure 4d represents the inverted binary image of the standardized grayscale image. If the melting is not appropriate, void spaces will occupy the melted region, which can cause porosity in fabricated parts. In the binary image (Figure 4c), distinct porous regions can be identified (shown by the black color) within the supposed melt area.

In the binary image (Figure 4c), distinct porous regions can be identified (shown by the black color) within the supposed melt area. Furthermore, Figure 4d represents a segmentation of the supposed melt area that is used for geometry measurements.In Figure 5a, the boundaries were detected in the binary image, where the boundaries of the object can be observed in red (outer boundaries of fabricated cylinders). The red boundaries represent the object shape and area, and, porosity can be observed as child boundaries with greenObject boundariesa)Child boundariesb)Figure 5. Porosity as child boundaries (a) and boundaries of objects (b)boundaries (boundaries within melt region). The small areas (porosity) covering less than 300 pixels were identified and later removed to obtain the overall size of the supposed melt area (Figure 5b). In this case, the binary image was converted to an inverted binary image and the small areas (porosity) were removed.

In this case, the binary image was converted to an inverted binary image and the small areas (porosity) were removed. The inverted binary image was inverted again to remove any residual areas within the image. In Figure 5b white areas represent the fabricated objects, and black areas represent the non-melting zone (powder bed) of each layer.No. of objects detected=6 Object No. 1 2 3 4 5 6Perimeter (pixel) Area (pixel) Temperature (K) Type 203 197 205 197 196 2033010 2802 3013 2822 2836 30231483 1481 1505 1502 1490 1501Circular Circular Circular Circular Circular CircularFigure 6. MATLAB output.Using a nested for loop and a conditional statement, the pixel values within each object boundary were read from the standardized image matrix, and all the pixel values were converted to temperature values. The average temperature values were calculated and printed as upshot in the MATLAB command window (Figure 6).

The average temperature values were calculated and printed as upshot in the MATLAB command window (Figure 6). Here, the area of detected object from the thermal image is based on a quantity of pixels detected where each pixel dimensionally represented a value of 170 microns [4]. A rounding error in the average temperature calculation (Figure 6) may be prevalent upon conversion. The maximum possible average temperature deviation was ~7.8K, depending on the round off error of each pixel value.The second technique compared was the use of edge detection, where the original image was converted to a grayscale image and edge detection was performed. Figure 7a represents the image after the detection of the edge in the grayscale image. In some cases, a discontinuous edge can be observed after edge detection. The edges were joined where the pixel of a discontinuity was near another pixel within a radius of four pixels.

In some cases, a discontinuous edge can be observed after edge detection. The edges were joined where the pixel of a discontinuity was near another pixel within a radius of four pixels. A defect of area in the fifth and the sixth object in the image (Figure 7b) can be seen that occurred due to discontinuous edges. Furthermore, the boundaries of the output layer were not precisely circular after the melting process, as shown in Figure 8a, which represents a graphical plot of the boundary for a singleObject 2Object 4Object 6a)b)Object 5 Figure 7. Edge detected in the image (a) and object after filling the edge area (b).Object 1Object 3b)a)c)Figure 8. Showing detected boundary of an object (a) having jagged edges that affect measurement, discontinuous edges (b), and unfilled boundary for an object (c).object. The jagged edges in Figure 8a can be explained by the system’s fabrication resolution or camera resolution.

The jagged edges in Figure 8a can be explained by the system’s fabrication resolution or camera resolution. In some cases, the discontinuous edges cannot be joined if the pixels on the edge do not meet within the radius of 4 pixels (Figure 8b). As a result, the object is incorrectly detected for discontinuous boundaries (Figure 8c). Because of the erroneous surface areas found and its consequence on average temperature values, the temperature measurement technique was omitted using the edge detection algorithm. Hence, the edge detection technique was not sufficient to check the precision of the output layer or compare with an input layer.Based on the algorithm mentioned earlier, the percentage error was found for each object of the selected layer after melting and compared to the ideal dimensions from CAD. Table 1 shows the data of error percentages for the area compared to the input layer of seven different thermal images corresponding to different fabrication layers.

Table 1 shows the data of error percentages for the area compared to the input layer of seven different thermal images corresponding to different fabrication layers. It was found that a different percent error of the fabricated area was found for each layer (always oversized from the CAD measurement). The range of percent error for object 5, for example, is 7.1% -16% and 5.4%-14% for object 6. The differences among fabricated areas between layers can be caused by corrugated surfaces of a built object, camera resolution, resolution of the fabrication system, and/or the material’s coefficient of thermal expansion (since images were taken during a heated environment).Table 1.

Error percentage of surface area pertaining to seven different images compared to their corresponding input layer images Object 2 error (%) 6.9 7.7 6.1 6.1 7.5 6.7 8.0Object 1 error (%) 16 17 16 16 17 16 16Object 3 error (%) 15 15 15 15 16 15 15Object 4 error (%) 6.5 7.1 5.2 6.0 7.1 5.8 7.7Object 5 error (%) 7.1 16 13 7.6 16 13 14Object 6 error (%) 14 8.1 5.4 15 8.2 5.9 9.2Layer 1 Layer 2 Layer 3 Layer 4 Layer 5 Layer 6 Layer 74.0 ConclusionThe methods presented in this research showcase the beginning steps of integrated metrology in advanced manufacturing systems and automatic monitoring of per-part thermal behavior and part quality. On the basis of analysis, thermal images provided sufficient data for the researchers to monitor and analyze the characteristics of EBM-fabricated objects. Two different methods for image analysis were compared, and it was determined that analysis on the basis of color was ideal for the images obtained in this research.

Two different methods for image analysis were compared, and it was determined that analysis on the basis of color was ideal for the images obtained in this research. Other methods of image analysis can be evaluated in future work. The geometrical data obtained in this study give a rough approximation as to the precision of the manufacturing process. Although a thermal camera was used in this study, higher resolution cameras can provide a better indication of fabrication precision on the basis of geometry. Furthermore, the geometric errors observed in this study can be attributed to various factors, including camera resolution, system resolutions, and the material’s coefficient of thermal expansion (since images were taken during a heated environment).

The MATLAB code generated in this study can be applied to other AM build processes, and further improvements of the software can be implemented that use the obtained data from the image for full spatial and temporal control as a means to obtain desired part mechanical properties. Furthermore, obtaining an indication of porosity was demonstrated in this study which can be used to perform parameter modifications that can help improve part quality and further lead to part qualification.AcknowledgementsThe research presented here was performed at The University of Texas at El Paso (UTEP) within the W.M. Keck Center for 3D Innovation (Keck Center), expanding recently to over 13,000 sq. ft. and providing access to state-of-the-art facilities and equipment as a result of funding, most recently, from the State of Texas Emerging Technology Fund. Support was provided, in part, by NIST grant 226351271A as well as the Mr. and Mrs. MacIntosh Murchison Chair I in Engineering Endowment.

tomating Sheet-Based Fabrication: The Conveyed-Adherent™ ProcessProf. Charles L. Thomas, Department of Mechanical Engineering, The University of Utah 3209 Merrill Engineering Building, Salt Lake City, Utah 84112 (801) 585-6939, fax: (801) 581-8692, e-mail: CThomas@ME.Mech.Utah.eduKenneth 1. Hayworth, Engineer, Ennex Fabrication Technologies 10911 Weyburn Avenue, Suite 332, Los Angeles, California 90024 (310) 824-8700, fax: (310) 824-5185, e-mail: KenH@Ennex.comCopyright © 1996, Ennex Corp.AbstractA new automated fabrication technology is described which breaks the fabrication process into spatially separatelayer-formation and layer-bonding stages. The technique uses sheet material on a substrate as feedstock and cuts cross-section contours into the material before conveying the material on the substrate to a stacking station.

The technique uses sheet material on a substrate as feedstock and cuts cross-section contours into the material before conveying the material on the substrate to a stacking station. Advantages include (a) speed, (b) versatility in fabrication materials, and (c) ability to fabricate hollows, embed or cast secondary materials, and selectively enhance or degrade material properties on a regional basis.A prototype fabricator has been built which automates all aspects of this process except weeding. Applications demonstrated using this machine include traditional rapid prototyping and visualization-model creation as well as functional machines taking advantage of embedding and cast-in-place techniques.I.IntroductionThe current commercial paradigm for automated fabrication (autofab) encompasses layered fabrication of plastic, paper, metal, or ceramic objects.

The paradigm consists of decomposition of a 3-D object into 2-D cross sections and sequential construction of each cross section in a "form in place" strategy. In a form-in-place technique, successive layers are bonded to the previous layer as they are formed. Each of the devices available for doing this has unique advantages and disadvantages which allow several commercial systems to compete successfully with no single system dominating all applications.This specialization is clearly demonstrated by comparing the Sanders thermoplastic droplet deposition process tothe laser cut paper layering process ofHelisys. The Sanders machine can produce layers as thin as 0.0004 inches which allows objects to be constructed with surfaces where the 3-D aliasing (or stair step effect) is greatly reduced. This process is excellent for producing small models ofjewelry or similar items.

This process is excellent for producing small models ofjewelry or similar items. The Helisys machine would have a difficult time producing these objects, and would require hand finishing to achieve a smooth surface fmish. However, the Helisys machine is capable of producing models up to 32" by 22" by 20" in a single build. The Helisys device is capable of building a model significantly faster than the Sanders machine[l]. Since the Helisys device only needs to laser-cut the edges of each cross section to create each layer, the process is very efficient for large heavy-bodied objects, but not as useful for thin-walled structures.To date, the prototyped objects are primarily considered models since their material properties and constructionaccuracy do not generally meet the capabilities of the intended production system[2,3].

In addition, while the construction speed varies from device to device, the current commercial systems typically cannot build at a rate much higher than 20 cubic inches per hour. The machine manufacturers recognize these limitations and are281introducing solutions to the problems. Many of the commercial systems are introducing techniques for producing patterns for investment casting or direct production of tooling for injection molding. Achievable tolerances remain in the range of 5 to 10 thousandths, however, which limits the use of these models in fine tolerance applications. 3D Systems has recently introduced a thermoplastic droplet deposition device that is an attempt to address the build speed issue. Their Actua™ device is reported to be less accurate than their stereolithography machine, but faster.In order to further address these problems, researchers are attempting to provide a larger variety of constructionmaterials.

They are attempting to work with metals and other traditional engineering materials[4,5,6]; they are attempting to work with multiple materials simultaneously[4,6]; and they are developing much faster processes aimed at production of large objects[7].II.Form-then-Bond FabricationThe work discussed in this paper is an attempt to improve the speed and flexibility of autofab processes by stepping outside the form-in-place paradigm to a "form-then-bond" technique. In this approach, layers of an object are first formed, and then registered and bonded together as separate operations. This change in paradigm produces devices that behave differently than the existing systems, with these differences resulting in both advantages and disadvantages. Physical realization of this process has been shown to be quite inexpensive.An inexpensive manual system based on the form-then-bond approach has given an indication of it's potential.

Physical realization of this process has been shown to be quite inexpensive.An inexpensive manual system based on the form-then-bond approach has given an indication of it's potential. lPSystem 5 Desk Top Rapid Prototyping offered by Schroff Development Corp. of Mission, Kansas, employs an off the-shelf cutting plotter along with custom modeling and slice generation software to create cross sections from sheet materials with a pressure-sensitive adhesive backing. The sheets are manually weeded of excess material and stacked after cutting. Although weeding and stacking each sheet can be tedious, construction of many objects can often proceed quite rapidly. A similar system was briefly offered several years ago by the Swedish company, Sparx AB. Case Western University and CAM-LEM Inc. have proposed an automated form-then-bond process with sheet handling performed by a robotic system.Automating a form-then-bond process produces added potential for improved build speed.

have proposed an automated form-then-bond process with sheet handling performed by a robotic system.Automating a form-then-bond process produces added potential for improved build speed. The form-in-placeparadigm requires that each layer be bonded and formed sequentially, in close succession. In a form-then-bond technique this is not required. It is not even necessary to perform both operations on the same machine. One of the anticipated implementations of this device includes one or more cutters that produce sheets ready for stacking, and one or more stacking devices that are fed a collated stack of construction sheets and produce the registered and bonded object. Using this arrangement, the two processes don't have to wait on each other.

Using this arrangement, the two processes don't have to wait on each other. If the stacker is faster than the cutter, one stacker could service several cutters, etc.This potential for build speed increase can be investigated by expressing the build time for a fabrication processin a general equation:nTbuild = T pre + L ~ayer( i) + Tposti=lwhere T pre is preprocessing time, T1ayer(i) is the creation time for the ith layer, and T pos1 is the post-processing time. While the details of what goes on during the pre- and post-processing times vary from process to process, we will focus here on the variation in T1ayer. The form this term of the equation takes depends on the specific technique used to create the layers of the object. We discuss here two overlapping classification schemes: 1) location offormation (form-in-place and form-then-bond) and 2) formation technique (volume building, periphery cutting, and 2-D mask).

We discuss here two overlapping classification schemes: 1) location offormation (form-in-place and form-then-bond) and 2) formation technique (volume building, periphery cutting, and 2-D mask). Only periphery cutting is today represented both in form-in-place devices, in which case it is called a "stack-first" process, or as a form-then-bond, "cut-first," process. For example stereolithography is a form-in-place volume building technique that creates a cross section by sequential curing voxel by voxel; Laminated Object Manufacturing is a form-in-place periphery-cutting (stack-first) process; Cubital's Solid Ground Curing is a form-in-place 2-D mask282process; and the JP5 (as well as the Conveyed-Adherent device discussed later in this paper) use a fonn-then-bond periphery-cutting (cut-ftrst) process.

Each layer creation scheme has a unique equation for calculating lIayer(i)' as shown in Table 1.Location of FonnationFonnation TechniqueExamplesLayer-creation timeVolume buildingSLA, SLS, FDMA. T;ayer(i) = ~ +_1 dvFonnin placeStack ftrstLOMT, layer (i)T Cl r +- V2D maskSolider (Cubital)Fonn then bondCut ftrstJP5, Conveyed-AdherentTable 1 Layer-creation time in various fabrication processes. In the equations, Tr represents the delay time between active geometry creation for each layer (e.g., recoating time for stereolithography or layer bonding time for LOM), Ai is the area of the ith cross section, Ci is the ith cut length, v is the cutting or laser-scanning velocity, d is the width of the line created by a volume-building process, and Tc is the cure time in a 2-D mask process. The second tenn on the right-hand side of each equation represents the time required for active geometry creation.

The second tenn on the right-hand side of each equation represents the time required for active geometry creation. Note that for a fonn-then-bond process the active geometry creation and the delay time (in this case layer registration and bonding) occur in parallel.Using the equations in Table 1 with constants selected for convenience (not exactly corresponding to any real process), the qualitative behavior of the different process classes can be investigated. The build speed as a function of geometric complexity (e.g., thin walled vs. thick walled) is investigated in Fig. 1 using a simple cube that is hollowed into a rectangular tube. The geometry is varied by changing the wall thickness of the tube, and the complexity is varied by adding multiple copies to the build volume. For this geometry the area of each layer is A= 10 - li2 , and the periphery length to be cut is C = 410 + 41i.

For this geometry the area of each layer is A= 10 - li2 , and the periphery length to be cut is C = 410 + 41i. Note that the build time increases for a volume-building process as the wall thickness increases, while the build time decreases for a periphery-cutting process (since the total18 . , . . . . - - - - - - - - - - - - - - - - - - - - - - - . . ,Volume building16en:; 12o ..c ~10 E E.:: 8 32 '3 6 c:o42 o ...1.-...=-~ -'10'Cut first- 'o0.511.522.5Wall Thickness (arbitrary units)Fig. 1 An example of how build time for the several types of fabricationprocesses can vary with respect to the geometry being fonned.2832cut length decreases with increasing wall thickness in this example). The 2-D mask process is independent of the wall thickness. The build time for the cut-first (form-then-bond periphery cutting) process is shifted downward since the cutting and stacking processes run in parallel.

The build time for the cut-first (form-then-bond periphery cutting) process is shifted downward since the cutting and stacking processes run in parallel. Note that the build time becomes independent of wall thickness when the periphery cutting time drops below the register and stack time.III. Automating Form-then-Bond FabricationEnnex Fabrication Technologies has automated form-then-bond fabrication with the development of theConveyed-Adherent™ process. This jump to full automation solves the difficulties of sheet handling and stacking and potentially provides significantly improved speed and convenience. Automation is achieved by the use of a substrate to carry each formed layer in succession to be stacked and bonded to the previous layers. This technique has the potential to address many concerns in the fabricator market and to open up new markets previously not in reach because of price. A schematic of the process is shown in Fig.

This technique has the potential to address many concerns in the fabricator market and to open up new markets previously not in reach because of price. A schematic of the process is shown in Fig. 2., ~Substrate-weeded material~New material laid down"SubstrateObject being fabricatedFig. 2 The two stages of the Conveyed-Adherent™ process, layer formation on the left and layer bonding on theright.Process OverviewThe basic construction material required by the Conveyed-Adherent process is any sheet material which can be cut and bonded to itself, along with a release liner or backing paper referred to here as the substrate. A 2-D plotting device is used to cut the outlines of cross sections of the desired object into the sheet material without cutting through the substrate. In addition to the outline contours, the plotter can also cut parting lines and outlines for support structures. When required, the sheet is then "weeded" to remove some or all of the negative material.

When required, the sheet is then "weeded" to remove some or all of the negative material. Finally the material is inverted so that the substrate is facing up and the material is brought into contact with the top of the growing object, and bonded to it. The substrate is then peeled off to reveal the new layer just added and a fresh surface ready to bond to the next layer. The process is started by applying an initial layer of adhesive on a stacking platform.The Prototype Genie ™ FabricatorThe Conveyed-Adherent process has been implemented in the laboratory of Ennex Fabrication Technologies in aworking prototype called the Genie fabricator, named for the mythical creature in Arab literature which could grant its owner any object desired. A picture of the machine, with a model of an automobile it has built placed in its build area, is shown in Fig. 3.284The prototype Genie fabricator operates automatically, except for weeding, with processing monitored by a variety of sensors.

3.284The prototype Genie fabricator operates automatically, except for weeding, with processing monitored by a variety of sensors. Material jams and other machine faults are detected by sensors for limits of travel and arrival of material at each station.The prototype Genie fabricator has built several objects, including various geometrical and automotive models inseveral colors. The models are attractive, although they suffer from problems due to registration errors and curl caused by internal stresses arising in the process. Research is underway to eliminate these problems and improve the quality of output of the fabricator. An example object made is shown in Fig. 4.,Fig. 3. The prototype Genie ™ fabricator in the laboratory of Ennex Fabrication Technologies.Fig. 4. Automobile model built on the prototype Genie fabricator at Ennex Fabrication Technologies. The car measures 9~ by 3314 inches by 3 inches high.

4. Automobile model built on the prototype Genie fabricator at Ennex Fabrication Technologies. The car measures 9~ by 3314 inches by 3 inches high. (Model data courtesy Viewpoint Datalabs International)MaterialsMany materials have been tested in the process, including label paper, adhesive-backed foam material, and signmaker's vinyl. Objects made from vinyl or paper are somewhat flexible due to the flowability of their pressure sensitive adhesive. The paper material can be made more rigid by coating with epoxy or glue. The foam material is used to build patterns for investment casting, resulting in metal objects. Ceramic and metal tapes for the process are under development by Lone Peak Engineering. Potentially any material can be used that can be produced as a sheet, cut by a knife, and undergo some form of bonding from layer to layer.PlottingLayer cross sections are cut using a cutting plotter designed for the sign making industry.

These plotters are capable of high dimensional accuracy with knife speeds up to 30 inches per second. Cutting plotters accept sheets of various widths up to over a meter wide and the vertical plotters, as opposed to bed plotters, have no limitation to length of sheet since the material is pulled back and forth through the machine. These machines can cut a wide variety of materials including the ceramic and metal tapes mentioned above, as well as very thin (less than a285thousandth on an inch) and thick (up to .035 inch and more) materials. The knives for these plotters include high precision blades;85 well as cryogenically treated metal blades to increase knife lifetime. A plotter is available which has a knife along with a color ink-jet printer. This would allow color to be incorporated into the material by a process proposed by Kmsie [8].

A plotter is available which has a knife along with a color ink-jet printer. This would allow color to be incorporated into the material by a process proposed by Kmsie [8]. In the prototype Genie fabricator, plotting is performed by a CAMM I cutting plotter made by Roland Digital Group.WeedingWeeding is the term used in the sign making industry for the process of removing the unwanted material around the letters and graphics coming from a sign plotter. To our knowledge there are no mechanical weeders available that can handle weeding of general complex shapes. These complex weeding operations are currently performed manually using tweezers and a good eye.A mechanical weeder allows one to overcome some of the limitations of stack-fIrst technology. The stack-fIrst process, in its current realizations, has no means of removing negative material during a build and hence produces output of a solid block of material.

The stack-fIrst process, in its current realizations, has no means of removing negative material during a build and hence produces output of a solid block of material. This requires that the negative material complementary to the object be "diced" by a cross-hatched pattern of parting lines, which add considerably to the process time. After fabrication, the negative material is chiseled away to reveal the fInished object. Thus, in the stack-fIrst process, enclosed voids cannot be made, as well as embedded macro-composite structures and colored objects.The engineering of a weeder, like the engineering of so many successful artifIcial intelligence programs, must usetricks and overt constraints to solve a less general problem. Some of the constraints are obvious, such as that the sheet is held down so as to remain a 2-D problem, and extra cuts can be provided to reduce diffIcult-to-weed shapes into smaller, simpler shapes.

Some of the constraints are obvious, such as that the sheet is held down so as to remain a 2-D problem, and extra cuts can be provided to reduce diffIcult-to-weed shapes into smaller, simpler shapes. Some "tricks" that can be used to reduce the amount of mechanical weeding necessary are discussed below.One trick to reduce the load on the weeder is to realize that most of the negative material can be dealt with using a few well-placed parting lines, so that the object can be easily broken out of its block-like enclosure by a simple turn of the wrist. The use of flexible materials makes this post-processing step particularly simple when compared to processes producing rigid fInal objects. A weeder is needed only in small regions of the object like enclosed voids or close approximations thereof.

A weeder is needed only in small regions of the object like enclosed voids or close approximations thereof. Therefore the weeder may be operating on only a fraction of the layers of an object and may be weeding only a small fraction of the negative space even on those layers.Another trick, called substrate weeding, has the potential to drastically reduce the work load of the weeder. Eachsuccessive layer of material containing the cross section of the desired object as well as its negative material is pressed on top of the growing object. If there is a void space below a section of the negative space (i.e. the last layer was successfully weeded) this section will have nothing to adhere to. It will remain on the substrate and be tossed out with it. Thus, in principle, all that is needed to prevent growth of one of the problem negative regions is to weed its first layer.

It will remain on the substrate and be tossed out with it. Thus, in principle, all that is needed to prevent growth of one of the problem negative regions is to weed its first layer. All subsequent sections will fInd no section below them to adhere to.The actual practice of this powerful technique is not as simple as the above description may imply. There areseveral factors that limit the ease with which substrate weeding can be used. One factor is overlap of negative material onto positive material. If a negative region grows larger with the height of the object, the section of negative material on any given layer will fInd at least a small ledge below it on which to adhere. The degree to which this causes a problem depends on how large the ledge is and on the relative strength of adhesion to the fabrication material and to the substrate. Another factor is cross-boundary linking.

Another factor is cross-boundary linking. Unlike a laser, a knife does not remove material when it makes a cut and the cut material has a tendency to self-heal. The linkage between already-cut sections within a layer is relatively weak, but can overpower lay-down adhesion for a small region whose cut periphery is large compared to its area.The prototype Genie fabricator does not have a mechanical weeder. Layers that require weeding are handled by aspecial diverter that routes material to a station where it can be weeded manually. Due to substrate weeding, the number of layers on which this must be performed is small.286Lay-Down, and Peel-OffThe cut and weeded sheet is fed to a lay-down platen. The platen aligns the new sheet with the growing object and applies the pressure needed to bond the layers. Several designs are possible, including the simple roller used in the Genie fabricator, as well as arc-sections of rollers with much gentler slopes.

Several designs are possible, including the simple roller used in the Genie fabricator, as well as arc-sections of rollers with much gentler slopes. A roller platen may approach the stack from the side or it may "land" on the stack from above. One problem that can occur is the formation of bubbles, trapped pockets of air, between the layers. Research at Ennex Fabrication has investigated the effect of various platen configurations on bubble formation and other relevant issues.After the material is laid down, the substrate must then be peeled off. For substrate weeding to occur there are two competing forces which must be balanced against each other. As described in the weeding section, one force is the adherence of the material to the last layer, and the other force is the adherence of the material to the substrate. The process must ensure that no positive material is carried away by accident and that no negative material needing weeding is accidentally laid down.

The process must ensure that no positive material is carried away by accident and that no negative material needing weeding is accidentally laid down. For instance, a small radius-of-curvature peel-off device will tend to favor leaving negative material.The prototype Genie fabricator uses a simplified design in which a single roller performs lay-down and peel-off simultaneously. The roller lays down the material on its leading edge and peels off the substrate at its trailing edge. The substrate, along with any substrate-weeded sections of the layer, are automatically dropped at the far end of the roller's travel. The roller "lands" on the stack by the raising of the stack under the control of a vertical sensor. Proper registration has been one of the difficulties to be overcome in the design of the fabricator. Table 2 documents the improvement in layer registration that has been realized through several design iterations.

Table 2 documents the improvement in layer registration that has been realized through several design iterations. Registration values reported are estimated values for registration error alone. They do not include error due to distortion, etc. The values in the first column are an estimate of typical registration error, ignoring the extreme values. Values in the second column represent the extreme values caused by machine jams, etc.Design iterationavg. error (in) max. error (in)frequency1st iteration0.0150.03550%2nd iteration0.0050.035<10%3rd iteration0.0020.020<1%Table 2 Registration improvement in the prototype Genie fabricator.

error (in)frequency1st iteration0.0150.03550%2nd iteration0.0050.035<10%3rd iteration0.0020.020<1%Table 2 Registration improvement in the prototype Genie fabricator. The average registration error of the prototype Genie fabricator, as well as magnitude and frequency of maximum registration error, has been reduced with each iteration of the design.Data FormatsThe Genie fabricator accepts a special layer-specification data format which includes the graphical instructions for each cross section and will include layer headers containing information about weeding, changing of feedstock ma~erial (such as color), and process pausing for manual embedding of components by the operator.A special feature of the Genie fabricator is the ability to change the length of the build envelope for any layer in abuild. This length is also specified in the data format.287IV.

This length is also specified in the data format.287IV. Demonstrated ApplicationsSeveral distinct advantages are realized in the form-then-bond paradigm. When using a cut-sheet process, it is possible to weed the negative areas of each construction layer before stacking, allowing the construction of hollow structures. These hollow structures can be filled during production producing cast in place internal structures. Objects may be inserted into these hollow spaces during construction producing a hybrid of fabricated and off-the shelf components. Construction sheets of different materials can be used resulting in objects built of several different materials in a single build. By combining the ability to embed components in voids and cast plastic into voids along with cut patterns designed to weaken the material selectively, the material properties within an object can be selectively specified.

One part of the object can be of rigid polyurethane, another bordering it can be of greater flexibility than even the original feed stock material. These techniques are demonstrated in the fabricated objects described below.Embedded and Cast-in-Place Car AxlesThe axles of two car models built in the Genie fabricator (e.g., see Fig. 4) serve as simple examples of cast-inplace and embedding techniques. Ifbuilt in the same vinyl used for the bodies ofthe cars, the axles would have sagged under the weight of the cars. Instead, voids were designed into the car bodies, wheels, and support material to allow casting plastic in a shape which would look like an axle in the visible portions and function like anchors inside the vinyl wheels and body. For one car, the axles were formed of polyurethane resin cast into the voids at the last layer before the void closed during the build.

For one car, the axles were formed of polyurethane resin cast into the voids at the last layer before the void closed during the build. For the other, steel axles were embedded in place with plastic extruded by a hot glue gun cast around them in the cavities. Both car models were in their complete form when the last layer was put down on the machine.Robotic ManipulatorsFig. 5 shows a dual-link robotic manipulator (or robot finger) built in the Genie fabricator with cast-in-place rigidplastic stiffeners, use of two materials to provide a rigid manipulator frame with flexible joints, and control cables inserted during fabrication. The manipulator is designed to be functional at completion of fabrication.The dual-link, four-degree-of-freedom manipulator demonstrates embedding of wire and tubing as well as casting-in-place of plastic. More importantly, the fmger's design necessitated a new way of thinking about how a fabricator is used.

More importantly, the fmger's design necessitated a new way of thinking about how a fabricator is used. Some of the shapes specified in the CAD file correspond to actual pieces of the build object made in the fabrication (vinyl sheet) material, while others correspond to voids and walls of molds to hold the casting resin, and still others must specify supports for embedded objects before the casting resin locks them in place. This isFig. 5 Dual-link robotic manipulator made on the prototype Genie fabricator at Ennex Fabrication Technologies.Left: CAD design; right: photograph of fabricated manipulator.288very different from designing a monolithic prototype model for a fabricator to build where only support structures need be specified. This has some of the flavor of designing a skyscraper where initial framing and access ways for build materials and construction equipment are just as important as the fmal product specifications.

This has some of the flavor of designing a skyscraper where initial framing and access ways for build materials and construction equipment are just as important as the fmal product specifications. This will become even more acute when more and more of the embedding and casting operation, here done by hand, are automated into the actual fabricator itself. Then the fabricator's ability to make arbitrarily complex monolithic shapes will provide problem constraints in the form of voids, supports, and fasteners to the very difficult problem of automated assembly itself.A CAD design, shown on the left side of Fig. 5, was created in TriSpectives, which specified the vinyl outer body of the three manipulator segments. A hollow cavity was specified inside each segment for the cast plastic "bones". A cylinder of 6 mm diameter was specified to run down the long axis of the manipulator to hold a flexible plastic tube.

A hollow cavity was specified inside each segment for the cast plastic "bones". A cylinder of 6 mm diameter was specified to run down the long axis of the manipulator to hold a flexible plastic tube. This tube and two springs fitting tightly around it hold the segments together, provide flexible joints, and in the final product provide a tube for electrical wiring, etc. Eight smaller-diameter cylinders were placed radially around this central tube to hold metal pull cables and their housings. Holes were left in the tops of the segment cavities to allow pouring of the casting resin. The resin seeps into the voids left in each segment and forms the bones while at the same time locking the vinyl, tube, pull wires, and joint springs in place. The pull-wire housing provides a low friction surface for the wire to run in even while the housing itself is embedded in the solidified resin.

The pull-wire housing provides a low friction surface for the wire to run in even while the housing itself is embedded in the solidified resin. The eight pull wire ends are terminated in the solidified resin locking them in place such that when one pulls the free end of the wire exiting the last segment of the finger the corresponding joint bends.The right side of Fig. 5 shows the fmished fmger mounted on a platform and connected to four servo motors allowing computer control of its four degrees of freedom. Future directions include embedding the servo motors, along with rechargeable batteries, micro controller, pressure and strain sensors, and associated wiring in a quadruped robot with legs resembling the above manipulator. From our experience with the above manipulator we expect this robot to be very durable.

From our experience with the above manipulator we expect this robot to be very durable. It will be completely protected from the elements since its sensitive components are encased in a homogenous mass of solidified resin and its movable joints have no bearing surfaces. When you include the fabricator's ability to provide arbitrary complexity to the shape of the outer body skin we see that this is a different sort of robot from the ones built in most research labs today.A single-link manipulator was built on the JP5 fabricator in a similar fashion. The frame for this manipulator was produced from multiple layers of paper; the hinge was constructed from four layers of vinyl. The dual-material build was realized by weeding an interior void from four interior layers of the paper and producing four vinyl construction sheets that contained only the hinge material.

The dual-material build was realized by weeding an interior void from four interior layers of the paper and producing four vinyl construction sheets that contained only the hinge material. To construct a multiple-material layer, first a paper layer was applied, then the vinyl layer was laid down in the voids of the paper layer.Motor-Driven Model CarA model car was fabricated with interior voids for embedded devices. In this example the solid STL file was modified by adding interior voids before construction. An electric motor, cables, battery pack, and switch were installed as the layers were assembled. After the last layer was in place, the switch was activated and the model car drove across the room.v.ConclusionsThe Conveyed-Adherent process is an automated "form-then-bond" fabrication technique that has been demonstrated in the production of engineering prototypes and functioning machines. The new process shows potential to be many times faster than existing devices.

The new process shows potential to be many times faster than existing devices. It has demonstrated fabrication in multiple materials side-by side. It also has demonstrated the potential to produce functioning machines through the use of embedded components and cast-in-place internal structures.289Notes and AcknowledgmentsThe authors would like to acknowledge the assistance of Marshall Bums, Paul Ashman, and Mark Walther ofEnnex Fabrication and Don Brock of the Manufacturing Processes Laboratory at the University of Utah. The Conveyed-Adherent process was developed by Ennex Fabrication Technologies and is covered by U.S. Patent Number 5,514,232 and other pending patents. The manual form-then-bond process commercialized by Schroff Development Corp. as lP System 5 Desk Top Rapid Prototyping was developed at the Manufacturing Processes Laboratory at the University of Utah and is protected by a pending patent.

as lP System 5 Desk Top Rapid Prototyping was developed at the Manufacturing Processes Laboratory at the University of Utah and is protected by a pending patent. "Conveyed-Adherent" and "Genie" are trademarks of Ennex Fabrication Technologies of Los Angeles.References1. Next-day delivery: A comparison ofrapid prototyping systems in Rapid Prototyping Report, pp. 4-6, July 1996.2. Bums, Marshall, Automated Fabrication-Improving Productivity in Manufacturing, PTR Prentice Hall, Englewood Cliffs, NJ, 1993.3. Thomas, Charles, An Introduction to Rapid Prototyping, Schroff Development Corp., Mission, KS, 1995.4. Merz, R., Prinz, F. B., Ramaswami, K., Terk, M:, and Weiss, L. E., Shape deposition manufacturing, Proceedings: SFF Symposium, pp. 1-8, 1994.5. Geiger, M., Steger, W., and Matthias, S., Multiphasejet solidification-A new process towards metal prototypes and a new data interface, Proceedings: SFF Symposium, pp. 9-16, 1994.6.

tomation of A Hybrid Manufacturing System Through Tight Integration of Software and Sensor Feedback J. K. Stroble*, R. G. Landers†, F. W. Liou* *Department of Manufacturing Engineering †Department of Mechanical and Aerospace Engineering University of Missouri-Rolla, Rolla, MO 65409 Reviewed, accepted September 14, 2006Abstract This paper presents a framework for the automation of the Laser Aided Manufacturing Process (LAMP) lab at the University of Missouri-Rolla. The groundwork for the proposed system involves the integration of the LabVIEW software package and a PXI-8195 real time controller with several sensors and actuators. The incorporation of all key control parameters into one virtual instrument will help achieve the goal of an automated hybrid system. To achieve this goal, a five-phase plan, which will be further discussed in the paper, has been developed.

To achieve this goal, a five-phase plan, which will be further discussed in the paper, has been developed. The first phase of this plan, which includes the deposition of a thin walled structure without DNC communication between LabVIEW and the CNC has been achieved, and will be the focus of this paper.IntroductionThe Laser Aided Manufacturing Process (LAMP) at the University of Missouri-Rolla (UMR) is a hybrid laser metal deposition (LMD) manufacturing system consisting of a laser, powder feeder, and motion system. The laser is used as a heat source while the powder feeder delivers metal powder at a specified rate into the path of the laser beam, thereby creating a melt pool. The laser beam and powder stream are directed vertically, while the substrate moves in three dimensions using the x, y, z, A, and B axes, molten tracks are deposited in layers, which cool rapidly to fabricate a part.

Sensors monitor the temperature, layer height, and melt pool geometry in real time via a real time (RT) control system.The overall goal of the UMR LAMP lab is the complete automation of the hybrid laser aided manufacturing process. To achieve this goal, a five-phase plan to automation has been developed. The five-phase plan involves utilizing sensor feedback to gain overall control of the diode laser, powder feeder, and motion system through a RT control system implemented on a single host computer. Virtual instruments (VI) created within the LabVIEW software package will be used to monitor, drive, and control the hybrid LMD process in real time.

Virtual instruments (VI) created within the LabVIEW software package will be used to monitor, drive, and control the hybrid LMD process in real time. The LabVIEW VI will include simulated controllers to compensate for undesired dynamics and noise, thus insuring accurate builds with a stable automated LMD process.The major focus of this paper will be to discuss the work performed to complete the first phase of the plan, which includes the deposition of a thin walled structure without DNC communication between LabVIEW and the CNC. To demonstrate the implementation of this phase, the paper will look at the equipment, software, and hardware required for control; the results from phase one’s implementation; and conclusions drawn from the first phase.Prior Work Hybrid manufacturing systems are a conglomeration of many off-the-shelf components that are combined in a modular fashion to achieve a new process.

Research was conducted on hybrid systems, individual components, and control applications. While the following research of586hybrid systems contains many of the pieces required for an LMD process, there is a general lacking in the areas of total system integration and control.Two pertinent real-time control applications dealing with laser and vision control are quality control inspection and position control. Real-time vision control for a fabric inspection system was shown to be very successful with dedicated hardware for the vision system being controlled via a Pentium 4 PC [1]. Morgan [2] developed a very reliable way of monitoring high power CO2 lasers based on the feedback of a light sensor and how to control the focal position of the laser.

Morgan [2] developed a very reliable way of monitoring high power CO2 lasers based on the feedback of a light sensor and how to control the focal position of the laser. Both applications mentioned are not associated with LMD, but contain aspects useful in the development of the hybrid system with relation to real-time control.Under the solid freeform fabrication (SFF) category, two articles by Malone [3,4] demonstrate successful types of positioning systems, deposition tools, and software. However, Malone has shown that small-scale systems are capable of deposition when being controlled by one computer system. Upton [5] has completed research on flexible manufacturing systems (FMS) where the key idea is that co-ordination of workflow is performed by a central control computer.

Upton [5] has completed research on flexible manufacturing systems (FMS) where the key idea is that co-ordination of workflow is performed by a central control computer. Both authors have laid groundwork in the area of hardware and software integration.Others at UMR have done research within the LAMP lab or dealing with lasers that is the most relevant to the automation of the LAMP lab. Specifically, Hua [6] has done extensive research in adaptive layer process control with lasers. Additionally, before the LAMP lab went through a major equipment upgrade in the summer of 2005, work went into system integration, experimental analysis, and modeling of the LAMP lab [7,8]. Although many of the components of the LAMP lab were changed, the fundamentals of the aforementioned research remains pertinent to the continued automation of the LAMP lab.FrameworkA five-phase framework has been proposed for the automation of a hybrid LMD system, which will be utilized in the UMR LAMP lab.

The framework lays out the major steps to achieving automation using real-time control hardware and integration of software with sensor feedback. Detailed steps for implementation of the five-phase framework are elucidated in the Methodology section. The parameters needed for successful framework completion are further discussed in the Parameters and Equipment for System Integration and Automation section.Phase 1: The first phase of automating the hybrid LMD process is to deposit a thin wall structure without DNC communications between LabVIEW and the CNC. Phase one demonstrates the ability to command the diode laser and powder feeder by the RT system and to simultaneously fabricate a part when a tool path is loaded on the CNC from another source.Phase 2: Phase 2 of the framework is similar to the first. A thin wall structure is deposited with DNC communication of the toolpath to the CNC from the VI running the laser and powder feeder.

A thin wall structure is deposited with DNC communication of the toolpath to the CNC from the VI running the laser and powder feeder. Depending on the type of CNC used and amount of on-board memory, drip- feeding of the tool path to the CNC may be required to fabricate the thin wall structure.Phase 3: Building upon the second phase, the third phase incorporates feedback from an intelligent vision system which monitors melt pool geometry. During deposition, the melt pool is monitored for elliptical geometry because as the substrate traverses, the round pool elongates. A feedback controller should be implemented that can interpret geometric feedback and compare it to the desired output. Once the melt pool leaves the allowed dimensions for the chosen laser power and powder mass flow rate, the deposition process reaches a warning mode.

Once the melt pool leaves the allowed dimensions for the chosen laser power and powder mass flow rate, the deposition process reaches a warning mode. If the vision system continues to report poor melt pool geometry for more than the allotted time, the LMD process faults and is shut down immediately587Phase 4: The fourth phase includes more sensor feedback by monitoring the temperature of the melt pool by a non-contact optical sensor. Due to the high priority of creating quality depositions, regulating the temperature of the melt pool is critical to achieving the desired microstructure. Modify the phase 3 controller to process additional data and simultaiously determine if the feedback is desirable. Once the measured temperature leaves the allowed range for the chosen laser power and powder mass flow rate, the deposition process reaches a warning mode.

Once the measured temperature leaves the allowed range for the chosen laser power and powder mass flow rate, the deposition process reaches a warning mode. If the temperature sensor continues to report an out of range temperature for more than the allotted time, the LMD process faults and is shut down immediately.Phase 5: The fifth phase incorporates the final sensor feedback, height of deposited layers, needed to complete the hybrid LMD system framework for automation. Incorporation of the laser displacement sensor feedback is an offline process that requires the deposition to pause so the sensor can scan the deposited structure, attain data, and display the data in real time. Modify the phase 4 controller to automatically process the offline feedback, and provide the option for an operator to decide if the data is acceptable.

Modify the phase 4 controller to automatically process the offline feedback, and provide the option for an operator to decide if the data is acceptable. If the data is acceptable, the LMD process will continue, otherwise it will be shut down.MethodologyDevelopment of the automation program to command and monitor a hybrid LMD system is comprised of several smaller tasks that build upon each other. The details needed to follow the proposed framework are contained within this section and describe the underlying work necessary for success. Completing the steps in sequence is critical when using this methodology.Step 1:Step 2:Step 3:Step 4:Step 5:Step 6:Test all LMD system devices for compatibility with the RT system hardware. Make necessary modifications to the devices as needed; such as building a special cable. Use the software package online diagnostic program to test if the software can the program accurately communicate with the devices.

Use the software package online diagnostic program to test if the software can the program accurately communicate with the devices. Measurement and Automation Explorer (MAX) is used for online diagnostic tests [9]. Create a basic VI to monitor the input and output of each device individually. The VI should contain at least a graph or chart that displays the output; fields for input parameters such as voltage, sampling rate, input channel, encoding type, ect., and a field to specify or monitor the save file path where the collected data will be stored. Perform open-loop step tests using the VI's created in Step 3 and record data to be analyzed. With a suitable mathematical software package, analyze collected data, and compare it to the predicted outcome. Look for system dynamics that will require additional modeling for compensation. Look for delays in the output that will inevitably affect the overall system performance.

Look for system dynamics that will require additional modeling for compensation. Look for delays in the output that will inevitably affect the overall system performance. If emulation is necessary, create mathematical models for the devices that exhibit significant dynamics to understand how to remove their disturbance from the overall system. Add code to the VIs created in Step 4 mimicking the mathematical models. Repeat Step 4. If emulation is not needed, then skip Step 5. For devices that only need to be monitored, new VIs will not be required in this step. Again, execute Step 4 using the VIs from Step 5 if emulation was used, until desired results are achieved. Develop an adequate controller that will regulate the output signals sent by the RT system to the controlled devices of the LMD process. A new VI should be created for each device and include the controller code. Execute Step 4 until the open-loop tests provide desirable results.

A new VI should be created for each device and include the controller code. Execute Step 4 until the open-loop tests provide desirable results. Next, update the new VIs toIf using LabVIEW,588589major role in the automation scheme of the RT control system for the hybrid LMD process. Therefore, a fast sampling controller, network card, analog and digital I/O ports, serial ports, hardware timers and counters, D/A converters, A/D converters, and hardware filters are some of the key aspects of a reliable RT control system. Conversely, a robust software package is required for overall tight system integration. LabVIEW, the software chosen for the LAMP lab, is a powerful software package developed by National Instruments. The LabVIEW software package is a robust and expandable software package for design, control, and testing [9]. Development of VI's, component control, and monitoring for the LAMP lab are completed as described in the methodology section.

Development of VI's, component control, and monitoring for the LAMP lab are completed as described in the methodology section. Figure 1 shows all the device inputs and outputs of the LAMP lab hybrid LMD process.The advantages to implementing an integrated system are three-fold. First, the hybrid LMD process can be made safer by becoming an automated process and removing people from directly interacting with the components and laser. Second, the options for control and feedback are endless and versatile. There are no limits on the number of VIs that can be created with the LabVIEW software package, so numerous programs can be developed and executed on the RT system or stored for later use. Thus, the hybrid LMD system is only limited by the hardware, which includes the I/O and CPU of the RT system. Third is repeatability leading to better quality control.

Thus, the hybrid LMD system is only limited by the hardware, which includes the I/O and CPU of the RT system. Third is repeatability leading to better quality control. With full automation, the hybrid LMD process will fabricate parts that have predictable and desirable characteristics more frequently.Some process parameters are not appropriate for real time control and should be held constant during the process of fabrication. The spot diameter provides the clad width and is determined by the focal length of the laser lens and the standoff distance. Thus, repositioning the z-axis can only change the spot diameter. This would require G codes to be sent to the CNC. Changes to the G and M codes sent to the CNC cannot be completed in real time because there is a delay when waiting for the last line of code in a program to be executed.

Changes to the G and M codes sent to the CNC cannot be completed in real time because there is a delay when waiting for the last line of code in a program to be executed. Another factor is that the setup of the powder feeder nozzle must ensure that the metal powder converges at the melt pool in a diameter roughly the size of the spot diameter. Altering the spot diameter would thus require an adjustment to the powder feeder nozzle, which cannot be done in-process. The table velocity is also not a candidate for real-time control. Only after a tool path program has been completed can the table velocity be changed because the whole program is sent to the CNC at once. Similarly, the tool path must also be set before the process begins.

Similarly, the tool path must also be set before the process begins. The two process variables that can be used for real time control are laser power and powder mass flow rate since they can be controlled independently of the other process parameters and the CNC.Key parameters for system integration are the ones that can be manipulated in real-time to induce a change in the final product or monitored for use with a feedback control scheme. By controlling and monitoring the key parameters, the quality of fabrication will increase and be repeatable. An overview of the parameters is given next along with how the device was affected by the steps presented in the methodology section.The main difficulty involved with controlling the powder mass flow rate in process is the natural delay that occurs between the control signal and the actual output. Powder mass flow rate is controlled by a command voltage, which regulates the rotational speed of the powder delivery shaft.

Powder mass flow rate is controlled by a command voltage, which regulates the rotational speed of the powder delivery shaft. The powder must then traverse the delivery system before entering the melt pool thereby creating a delay between the effective mass flow rate and the desired mass flow rate. Argon is used as the carrier gas for transporting the powder from the powder feeder to the laser collimator at a pressure of 40 psi. Also, a special cable was made to make the powder feeder mass flow rate (gpm) controllable by the RT system. Other considerations include the location590f where the powder stream converges to the location of the melt pool and preheating the powder to remove moisture. Preheating improves flow and helps minimize porosity in the finished part. Controlling the diode laser power by a command voltage was achieved by way of a special cable that connected the laser to the RT control system. The only delay is the 0.5 ms response time of the laser [11].

The only delay is the 0.5 ms response time of the laser [11]. The difficulty with controlling the laser power is determining what the desired laser power should be based upon the desired clad dimensions. Increasing the laser power increases the size of the melt pool and could increase the size of the deposition height if enough powder is present. The laser power must also be within a certain effective range for a given material since the final mechanical properties of the part, such as porosity, density, and microstructure, are closely related to laser power through melt pool temperature and solidification time. Laser power must also be large enough to induce melting in the substrate, but must also be below the point where dilution causes poor solidification.Real time monitoring of the melt pool length and width are important to maintain the dimensional accuracy during laser deposition [8]. Melt pool geometry is directly affected by the laser power and powder mass flow rate.

Melt pool geometry is directly affected by the laser power and powder mass flow rate. Dilution of the melt pool will result in poor cladding and produce unacceptable part quality. In order to monitor the melt pool geometry, a side bracket attached to the collimator emulating an axial mount with the use of two dichromic mirrors allows for a CMOS camera to acquire melt pool images during deposition in real-time. The length and width of the melt pool are extracted using an image-processing algorithm in real- time and used for feedback control during the last three phases of the framework.Layer height must be determined to calculate the number of layers that need to be run to minimize the use of raw material [8]. A non-contact laser displacement sensor is used to measure the layer height after an individual layer or a given number of layers have been deposited. Height is affected equally by the powder mass flow rate and the laser power.

Height is affected equally by the powder mass flow rate and the laser power. A higher laser power combined with more powder, leads to a bigger clad. In order to measure the height with the RT system reliably, a hardware filter was installed into the RT control system to alleviate most of the noise in the signal. The same is true for the temperature sensor, but with the addition of resistors to reduce the voltage output.Melt pool temperature is monitored continuously, in real time, using a dual-wavelength non-contact temperature sensor. If the temperature is too low, then the powder injected into the molten pool will not melt. Moreover, if the temperature is too high, it risks the danger of melting the previous layers too much or causing damage to the work piece [8].

Moreover, if the temperature is too high, it risks the danger of melting the previous layers too much or causing damage to the work piece [8]. The sensor measures the peak temperature of the melt pool formed during laser deposition and is used for feedback control during the last three phases of the framework.Direct Numerical Control (DNC) is a feature of the CNC machine that allows for a host PC with an RS-232 port to communicate with the CNC remotely. The 64Kb of memory local to the CNC is used when downloading a program at 9600 baud into the CNC memory for execution [12]. Since the CNC memory size is very small compared to a complete tool path program, the 64Kb of memory can then be used as a buffer for the program and frequently replenished by the remote PC until the full program has been loaded into memory and executed.

This is also known as “drip feeding.” The buffer fills after a few lines of code have been executed and continues to stay full at 256 lines of code until the last line of the program has been sent. However, the most important advantage to DNC is the way it handles large program files by drip feeding them to the CNC smoothly until the program is finished. This allows for large tool path programs to be automatically executed. Using the diagnostic software, it was discovered the CNC needed a special command to initiate DNC capabilities thus allowing for the phase two progress to begin.591ResultsPhase one of the LAMP lab framework has been completed and is demonstrated by the preliminary results shown in Figure 2. The thin wall structure was deposited semi-automatically, which means that the host PC communicating to the RT system commanding the laser power and powder mass flow rate did not drip feed the tool path to the CNC.

Another computer currently dedicated to performing DNC was used to send the tool path program to the CNC. Additionally, the main VI did not incorporate feedback control when the preliminary results were attained. The user of the main VI could control the powder feeder and laser voltage commands, and monitor and record their respective feedback signals. The integration of the software with the hardware was evident when the laser and powder feeder responded to the command signals without any complications, noticeable delay, or the loss of data samples. Given the robust nature of LabVIEW, the preliminary deposition task was simple to implement and was performed effortlessly by the RT system.Figure 2: Semi-Automatic Deposition of a 20 Layer Thin Wall StructureAs one can see from Figure 2, the deposition was very clean and had nice quality on the outside.

The first deposition (bottom) warmed the substrate, subsequently allowing the second deposition (top) to have better dimensional accuracy. Microstructure and porosity are still yet to be determined for the samples in Figure 2. To achieve such results, a powder mass flow rate of 8.25 gpm and a laser power of 700 W were used, which corresponds to a command voltage of 1.3 V and 6 V, respectfully.The correct command voltage for the laser and powder feeder were determined experimentally through open-loop step tests. Table 1 provides the steady state results of gpm and rpm for command voltages between 1–2 V, in 0.1 V increments. The rpm was recorded by the RT system at a sampling rate of 1000 Hz, and the caught grams of powder were measured on a scale. A VI was created to automatically send a command voltage to the powder feeder for one minute, shut off the powder flow by sending 0 V, and then stop the program.

A VI was created to automatically send a command voltage to the powder feeder for one minute, shut off the powder flow by sending 0 V, and then stop the program. During that minute, powder was captured in a glass jar at the end of the nozzle and weighed on a scale for 30 seconds to allow enough time for an approximate reading of total grams of powder, as recorded in Table 1.the four tests were averaged and checked for acceptable standard deviation. The results were suitable and can be found in Table 2. The data in Table 2 provides aConsequently,592reliable guide for the user when programming a VI for control, because the gpm has been correlated to command voltage. Figure 3 shows the relationship between command voltage and the powder mass flow rate with a calculated slope of 10.4 when analyzed using the least squares method. Pleasingly, the correlation coefficient was found to be 0.999.

Pleasingly, the correlation coefficient was found to be 0.999. Deviation within the rpm test data is negligible in most cases, but the gpm deviation was large for voltages of 1.00, 1.60, 1.70, and 2.00. It is hypothesized that fluctuations between gpm test results are mainly caused by the powder wheel mechanism consisting of a cam and flexible follower within the powder feeder. The position where the powder wheel starts and stops during each test has a great impact on the amount of powder released by the mechanism, because each cycle of the powder wheel is not identical. Large deviations were also partially due to measuring the grams by hand with a scale and recording the value that was displayed most frequently within the 30 seconds the jar rested on the scale.

Moreover, the type of distribution system installed before the collimator splits the main powder stream into four, and can become clogged, statically charged, or leak carrier gas, which can deteriorate powder delivery performance significantly.Table 1: Results of Powder Mass Flow Rate Open Loop TestsCommand Voltage (V)Recorded gpm (approx.)RPM Recordedgpm (approx.)RPM Recordedgpm (approx.)RPM Recordedgpm (approx.)RPMtest 1test 2test 3test 41.00 1.104.90 6.020.5047 0.65484.60 5.800.499 0.64974.80 5.900.5024 0.65754.73 5.920.5013 0.65721.20 1.307.15 8.420.811 0.95647.13 8.330.8067 0.95627.14 8.230.8089 0.95577.19 8.250.8071 0.95721.40 1.509.50 10.401.115 1.2639.37 10.381.11 1.2649.40 10.401.11 1.2639.36 10.351.11 1.2621.60 1.7011.48 12.521.422 1.57811.52 12.221.422 1.5811.39 12.301.425 1.58111.30 12.231.427 1.581.80 1.9013.00 14.351.737 1.8813.13 14.461.739 1.88113.00 14.291.742 1.88413.13 14.251.727 1.8822.0015.142.04415.032.04715.412.04715.842.047Table 2: Averages and Standard Deviations for Data in Table 1Command Voltage (V) GPM Avg.

RPM Avg. GPM Std. Dev. RPM Std. Dev.1.00 1.10 1.20 1.30 1.40 1.50 1.60 1.70 1.80 1.90 2.004.76 5.91 7.15 8.31 9.41 10.38 11.42 12.32 13.07 14.34 15.360.5019 0.6548 0.8084 0.9564 1.1113 1.2630 1.4240 1.5798 1.7363 1.8818 2.04630.13 0.09 0.03 0.09 0.06 0.02 0.10 0.14 0.08 0.09 0.360.0024 0.0036 0.0020 0.0006 0.0025 0.0008 0.0024 0.0013 0.0065 0.0017 0.0024593Figure 3: GPM Test Results of the Remotely Commanded Powder FeederFigure 4: RPM Test Results of the Remotely Commanded Powder FeederFigure 4 relates the average command voltage to the rpm. When the rpm data was analyzed using the least squares method, the slope was found to be 1.5. The correlation coefficient was found to be exactly 1.000 indicating a nice linear relationship as shown in Figure 4. Finally, the gpm and rpm test results were correlated in Figure 5 and the slope was found to be 6.7 by the least squares method.

Finally, the gpm and rpm test results were correlated in Figure 5 and the slope was found to be 6.7 by the least squares method. The results in Figure 5 were greatly affected due to the powder feeder mechanism and powder distribution system as previously mentioned. However, the relationship between the rpm and gpm is approximately linear with a calculated correlation coefficient of 0.998. Deposition test results have proven the collected data in the voltage range of 1-2 V to be reliable for use with the LMD process.Correlation between the commanded voltage and output wattage to the substrate was conducted using a Coherent Power Meter with the water-cooled LM5000 sensor head, rated for 5 kW. The sensor head was placed below the collimator at a standoff distance of 14.478 mm (0.57 in), and a voltage was commanded in 1 V increments to the laser by the laser VI. The bolded columns of Table 3 list the given documentation of the diode laser.

The bolded columns of Table 3 list the given documentation of the diode laser. The recorded measurements from the power meter tests at the substrate are labeled Pm Test, and the data standard deviation are in Table 3. Correlation between the provided documentation and the power measured at the substrate is in Figure 6. By the least squares method, the slope for the given information was found to be 167.00, and the slope for the measured information was found to be 128.00. It was calculated that the laser output correlation coefficient of the Pm Test average was 0.990, which is demonstrated by the large deviations at 1 and 10 V, where as, the given information correlation coefficient was 0.995. Due to losses in heat and the fiber optic medium, the power meter displayed a lower output wattage than what was to be expected as per the diode laser documentation. Furthermore, at the lower range of the voltage input, the output wattage is very close to the provided documentation.

Furthermore, at the lower range of the voltage input, the output wattage is very close to the provided documentation. It is only at higher command voltages that the laser does not perform as expected.594Table 3 : Laser Power Meter Test Results Vc (V)Pm Test 1 (W)Pm Test 2 (W)Amps Displayed (A)Pm Test 3 (W)Pm Test 4 (W)Amps Displayed (A)Pm Std.

Dev.Nuvonyx Displayed Amps (A)1 2 3 4 5 6 7 8 9 100 17 220 380 560 700 820 940 1030 10500 17 220 380 550 700 840 930 1020 10205 11 16 21.5 27 32.5 38.5 43.5 49 49.50 17 220 370 520 660 800 900 1010 10500 17 220 370 510 650 770 890 990 10205 11 16 21.5 27 32.5 38.5 43.5 49.5 49.50 0 05.7735 23.8048 26.2996 29.8608 23.8048 17.0783 17.320510 15 20 25 30 35 40 45Figure 5: GPM Results Compared to RPM Results of the Remotely Commanded Powder Feeder595Nuvonyx Output Power (W)57 225 399 574 737 884 1014 1124Figure 6: Comparison of Actual Laser Output to Given Laser DocumentationConclusions and Future WorkThe framework for accomplishing the goal of automating the UMR hybrid LMD system has been presented. By following the presented methodology for integrating hardware and software, individual manipulation and monitoring of laboratory components has been achieved successfully.

By following the presented methodology for integrating hardware and software, individual manipulation and monitoring of laboratory components has been achieved successfully. Methodology steps one and two proved to be very helpful in alleviating many unseen problems that did not seem evident in the beginning. Mainly, the temperature sensor needed to be modified for use with the RT system. Preliminary results were demonstrated through deposition samples as shown in Figure 2. The collected data presented in the results section demonstrates that phase one of the framework was successfully completed, because the main VI was only given control parameters and did not rely on feedback.

Integration of the software package, RT system, and LMD components was confirmed to be imperative and achievable for the success of full automation.The future work needed for completing the framework is to actively send information from the RT system directly to the CNC by way of RS232 communication to complete the DNC requirement of phase two. Once the DNC is completed, the last three phases will incorporate the feedback of the monitoring devices and how they interact with the overall system. A robust controller will need to be developed that can handle the feedback from three devices adequately. Real-time processing of feedback from devices simultaneously and driving the computed error signal to a minimum will be the capabilities of the controller.

Real-time processing of feedback from devices simultaneously and driving the computed error signal to a minimum will be the capabilities of the controller. After feedback control is in working order, fault conditions will be added to increase the quality of deposited parts created in the LAMP lab.596Acknowledgments This research was supported by the National Science Foundation Grant Number DMI-9871185, the grant from the U.S. Air Force Research Laboratory contract # FA8650-04-C-5704, and UMR Intelligent Systems Center. Special thanks to Robert Nagel for his assistance with this paper. Their support is greatly appreciated.References Cho, C., Chung, B., Park, M., “Development of Real-Time Vision-Based Fabric 1. Inspection System,” IEEE Transactions on Industrial Electronics (2005), Vol. 52, No. 4, pp. 1073-1078. 2. “Real-Time Process Control CO2 Laser Welding and Direct Casting: Focus and Temperature,” ICALEO Conference (1997), pp. 290-299. 3.

Specifically, Gaussian process regression is applied to formulate a surrogate model between the critical process parameters (i.e., laser power, scanning speed) and the residual porosity of the nitinol samples. Then Bayesian optimization is integrated to successively explore the design space to search for the optimal process parameters. These two methods are integrated to find the global optimum iteratively. Compared with the traditional trial-and-error methods, the proposed method can quickly find the optimal process parameter for the high-quality nitinol samples, especially with many process parameters, and accelerate the innovations with nitinol in additive manufacturing.Keywords: Nitinol, Laser beam-powder bed fusion, Process parameters, Bayesian optimization1 IntroductionNitinol (nickel-titanium alloy) typifies a generic class of shape memory alloys ranging in composition from 52 to 56 weight percent nickel and balance titanium [1].

They could undergo a large deformation (10-30 times as much as ordinary metal), remain in the deformed shapes when external forces being removed, and instantaneously restore to original shapes after being heated over transformation temperatures (known as “pseudoelasticity”) [1-5]. Moreover, nitinol has good biocompatibility to the human body [6-8] and has wide applications in medical devices, ranging from orthodontic, archwires, endoscopic instruments to endovascular stents [9]. Concurrently, the global nitinol medical devices market, evaluated at $14.52 billion in 2018, is projected to expand at a compound annual growth rate of 8.5% from 2019 to 2027 due to increasing demands for minimally invasive procedures [10].The laser beam-powder bed fusion (LB-PBF) process is a promising additive manufacturing (AM) technology to provide customized nitinol medical products with reduced production ramp-up time and design customization [11].

This process has the following advantages: (1) No additional tool is needed when prototyping different products; (2) Unused284material could be recycled to reduce the overall material costs; (3) The product design is customizable through the computer-aided design.However, researchers found that inappropriate fabrication conditions lowered the quality of nitinol products, which obstructed the applications of the LB-PBF process on nitinol [12-18]. More specifically, improper process parameters caused defects, cracks, and inferior surface quality (J/mm3) has been defined in of the LB-PBF fabricated nitinol products. An energy density Equation (1) with multiple process parameters in conjunction [19]: 𝐸𝐸𝑃𝑃 𝑣𝑣 × ℎ × 𝑡𝑡is layer where thickness (mm). Mahmoudi et al. [12] observed high porosity experimental samples with visible cracks when the energy density was lower than 30J/mm3. Walker et al.

Mahmoudi et al. [12] observed high porosity experimental samples with visible cracks when the energy density was lower than 30J/mm3. Walker et al. [17] detected an increased number of defects on samples produced with constant laser power and rising scanning speed due to the low energy density. Contrarily, they observed that the high energy density, comprised of high laser power and low scanning speed, increased melt pool dynamics to negatively affect the powder deposition in successive layers, leading to low-density samples. Wang et al. [16] found large amounts of keyholes and gas pores in the sample fabricated with 120W laser power and 400mm/s scanning speed. The numbers of keyholes were reduced remarkably with a 900mm/s scanning speed in their follow-up experiments, indicating the high energy density was the primary reason.

The numbers of keyholes were reduced remarkably with a 900mm/s scanning speed in their follow-up experiments, indicating the high energy density was the primary reason. These defects, pores, and surface cracks due to unsuitable process parameters could be the main sites for crack initiation, which negatively affected the fatigue behavior of the nitinol products [20, 21].is laser power (W),is scanning speed (mm/s),is hatch distancing (mm), and𝐸𝐸 =𝑃𝑃𝑣𝑣ℎMoreover, some inappropriate fabrication conditions could cause the chemical impurities of the nitinol products. Since nickel is more volatile than titanium at elevated temperatures, several works [16-18, 22] pointed out that nickel evaporated during the LB-PBF process when the energy density reached a higher value above 200J/mm3. Biffi et al. [18] found that the nickel component decreased from 51% to 50.4% when the energy density increased from 63J/mm3 to 300J/mm3.

Biffi et al. [18] found that the nickel component decreased from 51% to 50.4% when the energy density increased from 63J/mm3 to 300J/mm3. Some research [22-24] indicated that a reduction of 0.85% of nickel content led to a 30ºC increase in the transformation temperature of nitinol. Therefore, the nickel evaporation during the LB-PBF process due to high energy density could handicap the pseudoelasticity property of the nitinol and their applications of thermal-activated products.Since the relative density could reflect porosity levels and chemical impurities, many researchers searched for the process parameters of the LB-PBF process that produced high-density nitinol samples. In many works, experiments comprising different process parameters were conducted to find the optimal process parameters. Biffi et al. [18] ran 42 experiments involving seven unevenly distributed levels of laser power ranging from 30W to 175W and six levels of scanning speed from 500mm/s to 3000mm/s.

Biffi et al. [18] ran 42 experiments involving seven unevenly distributed levels of laser power ranging from 30W to 175W and six levels of scanning speed from 500mm/s to 3000mm/s. The sample fabricated with 125W laser power and 1250mm/s scanning speed reached the highest relative density above 99%. Walker et al. [17] conducted 48 experiments to study the relationship between process parameters (i.e., laser power and scanning speed) and relative density of LB-PBF fabricated nitinol samples. Their experiments covered ranges of laser power (50W to 300W) and scanning speed (250mm/s to 2000mm/s). They used 250W laser power and 1250mm/s to fabricate the sample with the highest relative density (> 98%). Wang et al. [16] studied the effect of varying single process parameter on the relative density285𝑡𝑡(1)of their nitinol samples.

Wang et al. [16] studied the effect of varying single process parameter on the relative density285𝑡𝑡(1)of their nitinol samples. They set different levels for scanning speed, hatch distancing, and laser power and conducted 22 experiments in total to fabricate high-density samples (>99%) with an energy density of 41J/mm3. Zhu et al. [15] fabricated nine nitinol samples under different conditions to build an analytical model based on the Eagle-Tsai simulation model. They used the analytical model to find an area of fabrication conditions free of defects such as keyholes, lack of fusion, and balling. They produced fully dense nitinol sample (>99%) with the process parameters in the defect-free area.Besides experiments and simulations, several machine learning methods have been applied to improve the quality of LB-PBF fabricated nitinol products. Mehrpouya et al.

Mehrpouya et al. [25, 26] used the artificial neural network (ANN) algorithm to predict functional properties of nitinol samples such as strain recovery ratio and transformation temperatures based on process parameters (i.e., laser power, scanning speed, and hatch distancing). Their ANN models achieved high accuracy (R2 values above 0.97) on the test data. Mahmoudi et al. [12] utilized the linear discriminant analysis to predict the occurrences of surface cracks depending on the linear energy density (laser power/scanning speed) values. The accuracy of their model was above 90%.

The accuracy of their model was above 90%. These models could accurately predict the functional properties of LB-PBF fabricated nitinol samples based on the process parameters, but they lack the capability to identify the optimal process parameters.In summary, there are two primary limitations among the state-of-the-art fabrication of the nitinol products via the LB-PBF process: (1) Inappropriate process parameters result in the low relative density of the nitinol products, which cripples their functions and future applications; (2) The related studies conducted time-consuming and costly experiments to test the effects of different process parameters on the relative density. In this case, we propose a nonparametric Bayesian framework to address these limitations by systematically combining the fabrication of nitinol samples through the LB-PBF process with the machine learning and optimization procedure in a physical-based data-driven manner.

The framework enables an efficient exploration within the design space of process parameters to find the optimal one that leads to the fabrication of high- density nitinol samples. Therefore, it could further unleash the potential of the LB-PBF process in nitinol applications.2 MethodologyThe objective of the Bayesian framework in the present work is to efficiently find the optimal process parameters (i.e., laser power and scanning speed) to fabricate high-density nitinol samples through the LB-PBF process. As shown in Figure 1, the proposed Bayesian framework encompasses developing the initial input-output dataset, applying the iterative process, and acquiring the optimal process parameters. The residual porosity (100% - relative density) values are used as the output in the initial dataset.

The residual porosity (100% - relative density) values are used as the output in the initial dataset. These data enter the iterative process, which consists of the following steps: (1) A Gaussian process (GP)-based surrogate model builds a relationship between process parameters and relative residual porosity through prediction; (2) The Bayesian optimization explores the surrogate model to search for the optimal process parameters; (3) The Bayesian update adds data from new fabrications to the dataset and iterates from the first step. The iterative process ends when the Bayesian optimization results converge, which leads to the acquisition of the optimal process parameter.286Figure 1. The steps of the proposed Bayesian framework to find the optimal process parameter for LB-PBF fabricated nitinol samples.

The steps of the proposed Bayesian framework to find the optimal process parameter for LB-PBF fabricated nitinol samples. The iterative process (red arrow) will be ceased when the Bayesian optimization results converge.2.1 Gaussian process (GP)-based surrogate modelFor the iterative process, we first utilize Gaussian process regression (GPR) to construct the surrogate model for mapping the nonlinear relationship between the process parameters and the residual porosity based on the initial input-output dataset. The GPR has been widely used to build computationally inexpensive (cost less than 1 minute) and high-accuracy (low mean absolute prediction error) surrogate models for process optimization in additive manufacturing [27-29].

Hence, we use the GPR to build the surrogate model for the Bayesian optimization to search for the optimal fabrication condition.The GPR predefines a prior Gaussian distribution with mean μ and covariance C over theregression functionwithout a parametric form:𝑓𝑓Our task is to utilize the function(1) to build the relationship between the process parametersand residual porosity:𝑓𝑓 ~ 𝐺𝐺𝑃𝑃(𝜇𝜇, 𝐶𝐶)𝑓𝑓𝒙𝒙𝑌𝑌(2)as the fabrication conditions of the nitinol is the random noise [27]. As a samples, which consist of the laser power and scanning speed. 𝒙𝒙 furnishes a notion of similarity between the fabrication conditions using its covariance prior, function (e.g., exponential kernel function).

𝒙𝒙 furnishes a notion of similarity between the fabrication conditions using its covariance prior, function (e.g., exponential kernel function). Therefore, closer fabrication conditions lead to 𝐺𝐺𝑃𝑃 more similar residual porosity values, enhancing the predictive power of the GP-based surrogate model.In the present work, we assign the predictors 𝑌𝑌 = 𝑓𝑓(𝒙𝒙) + 𝜖𝜖𝜖𝜖𝐶𝐶(·,·)287For modeling the fabrication conditions and residual porosity values, we establish training and the random variablesdata comprising of various fabrication conditionsof residual porosity values following a multivariate Gaussian distribution:(𝒙𝒙1, 𝒙𝒙2, … , 𝒙𝒙𝑛𝑛)𝒇𝒇𝑡𝑡𝑡𝑡𝑡𝑡𝑡𝑡𝑛𝑛𝑡𝑡𝑛𝑛𝑡𝑡(3)where meanis the expected porosity of the nitinol samples; the covariance function𝒇𝒇𝑡𝑡𝑡𝑡𝑡𝑡𝑡𝑡𝑛𝑛𝑡𝑡𝑛𝑛𝑡𝑡 ~ 𝒩𝒩�𝝁𝝁, 𝜮𝜮𝑡𝑡𝑡𝑡𝑡𝑡𝑡𝑡𝑛𝑛𝑡𝑡𝑛𝑛𝑡𝑡�encodes the correlations among the porosity variablesbased on the similarity 𝜮𝜮𝑡𝑡𝑡𝑡𝑡𝑡𝑡𝑡𝑛𝑛𝑡𝑡𝑛𝑛𝑡𝑡 =.

An acquisition function is used to direct sampling to areas where an improvementin Eq. (2) of the GPR. Bayesian optimization incorporates prior belief aboutprocess 𝒇𝒇 approximates over the current best observation is likely to occur.𝒇𝒇𝐺𝐺𝑃𝑃(𝜇𝜇, 𝐶𝐶) 𝒇𝒇Bayesian optimization maximizes acquisition function values based on the surrogate model . The expected improvement (EI) is adopted as the acquisitionto find the next sampling point function in this work, which is defined as the following equations: 𝒙𝒙′′(6)𝒙𝒙= argmax𝒙𝒙𝐸𝐸𝐸𝐸( 𝒙𝒙)288+is the lowest residual porosity values of the nitinol samples so far under the where . According to its definition, the EI acquisition function evaluates the fabrication condition + expected amount of improvement in the objective function.

According to its definition, the EI acquisition function evaluates the fabrication condition + expected amount of improvement in the objective function. It is often maximized with the implementation of Newton’s Method.EI(𝒙𝒙) = 𝔼𝔼 max(𝑓𝑓(𝒙𝒙) − 𝑓𝑓(𝒙𝒙), 0)𝑓𝑓(𝒙𝒙)+𝒙𝒙has been incorporated into Bayesian optimization to prevent premature convergence by overexploiting an area. To illustrate the overexploiting, we assume be the posterior standard deviation of the additive noise. The Bayesian optimization, after each iteration 𝜎𝜎𝐹𝐹(𝒙𝒙) satisfies the following of the acquisition function, evaluates whether the next condition equation:A high exploration ratio𝑡𝑡𝜎𝜎is the standard deviation of the posterior objective function at condition. Let𝒙𝒙𝒙𝒙′′In consequence, the algorithm declares that area near condition is overexploiting. ) < 𝑡𝑡𝜎𝜎𝜎𝜎 Correspondingly, the acquisition function modifies the kernel function by multiplying kernel ) with iteration numbers [31].

The predicted optimal from an unexplored area tends to have a high variance, indicating high uncertainties. Hence, we fabricate new samples using the predicted optimal fabrication condition and collect new data to reduce the uncertainties of these unexplored areas in the surrogate model. The Bayesian optimization algorithm will explore the updated surrogate model to search for the optimal conditions for further new fabrications. This process is iterated until the optimization results converge, which indicates that the global optimum has been found.There are several benefits for applying the developed Bayesian framework to the process optimization of LB-PBF fabricated nitinol samples. First, it reduces the number of experiments in finding optimal fabrication conditions. Compared to the full factorial design, which runs all combinations of process parameters, this Bayesian framework could work with fewer fabrications.

Compared to the full factorial design, which runs all combinations of process parameters, this Bayesian framework could work with fewer fabrications. For instance, a full factorial design with five levels of laser power and scanning speed requires twenty-five experiments in total. The full factorial design could result in exponentially increased experiments if new factors are added. In contrast, in the case study, the Bayesian framework starts with nine fabrications and needs approximately five iterations to find the optimal fabrication conditions, significantly improving the efficiency and reducing the experimental cost. As the framework concentrates on the areas which possibly include the optimal condition based on the289𝜎𝜎(7)Bayesian optimization algorithm, there is no need to cover all combinations of process parameters through costly experiments. More importantly, the global optimal fabrication condition could be found through this framework.

More importantly, the global optimal fabrication condition could be found through this framework. In contrast to finding the lowest residual porosity values from the surrogate model through interpolation, the EI acquisition function with a high exploration ratio could systematically explore the design space and prevent premature convergence of the 𝑡𝑡𝜎𝜎 optimization results.In summary, the proposed Bayesian framework could find the optimal fabrication conditions for the LB-PBF process to produce high-density nitinol samples with a reduced number of fabrications.

It could be easily extended to process parameter optimization in the LB-PBF process based on other required properties (i.e., transformation temperatures, elongation, tensile strength) for various applications, which entitles the proposed Bayesian framework great potential to improve the functional properties of the LB-PBF fabricated nitinol.3 Case studyThe proposed Bayesian framework is applied to a real-world case study to find the optimal process parameter of the LB-PBF process that fabricates high-density nitinol samples. In the case study, the nitinol samples are fabricated as cubic blocks. We fabricate these blocks through the LB-PBF process with different process parameters and correspondingly extract the residual porosity values to build the initial input-output dataset. Afterward, we conduct the iterative process of the proposed Bayesian framework with two more fabrications based on the Bayesian optimization result at each iteration.

Afterward, we conduct the iterative process of the proposed Bayesian framework with two more fabrications based on the Bayesian optimization result at each iteration. The case study result demonstrates the feasibility and effectiveness of the proposed Bayesian framework.3.1 Data collectionAn initial batch of 9 nitinol blocks (12.7mm×12.7mm×20mm) has been fabricated through the LB-PBF process with different fabrication conditions. These conditions cover a range of 100W to 200W of laser power and 1000mm/s to 1400mm/s of scanning speed. Since this work focus on the effects of the laser power and scanning speed on the residual porosity, the hatch distancing, and layer thickness are kept as 120 µm and 30 µm, respectively, throughout the whole process.Figure 2 and Figure 3 display the finished nitinol blocks and their corresponding microstructure images. The microstructure images in Figure 3 demonstrate the effects of process parameters on the microstructure of the nitinol blocks.

The microstructure images in Figure 3 demonstrate the effects of process parameters on the microstructure of the nitinol blocks. The blocks fabricated with an energy density lower than 30J/mm3 (e.g., blocks 1.1, 1.4, 1.7, and 1.8) have inferior qualities with a large number of defects shown as the dark regions.290Figure 2. Images showing the finished nitinol blocks through LB-PBF under different fabrication conditions.Figure 3. Microstructures as sectional views at X-Y direction of the nitinol blocks in the ascending order of energy density. The images show that blocks 1.7, 1.4, 1.1, and 1.8, fabricated with lower energy density values, have large amounts of defects shown as black regions.The proportions of the defect region are quantified as the residual porosity values. These values are extracted from the microstructure images, binarized to separate the defects from normal printings.

These values are extracted from the microstructure images, binarized to separate the defects from normal printings. The observed residual porosity values of the nitinol samples and their fabrication conditions are summarized in Table 1. Several blocks (e.g., blocks 1.3, 1.5, and 1.6) have more values since they obtain more microstructure images.Table 1. The fabrication conditions and corresponding residual porosity of the nitinol blocks.

According to the prediction surface shown in Figure 4(a), the area where the laser power is from 140~200W and scanning speed from 1000~1200mm/s has the lowest predicted mean residual porosity values (<5%). However, it is noticeable that the predicted mean residual porosity values within this area have a standard deviation from 4~10% based on the estimated standard deviation shown in Figure 4(b). The lowest predicted mean residual porosity value of 0.78% is detected with a 4.3% standard deviation at the first iteration of the Bayesian optimization. The corresponding laser power is 150W, and the scanning speed is 1070mm/s.Figure 4. (a) The prediction surface of mean residual porosity from the surrogate model at different fabrication conditions. (b) The estimated standard deviation at different fabrication conditions from the surrogate model.Figure 5. The first iteration of the Bayesian optimization with 50 evaluations to find the optimal fabrication condition.

The first iteration of the Bayesian optimization with 50 evaluations to find the optimal fabrication condition. The lowest predicted mean residual porosity is 0.78% with 150W laser power and 1070mm/s scanning speed.292The results of the Bayesian optimization in terms of predicted optimal residual porosity values and their corresponding process parameters at each iteration are summarized in Table 2, comparing to the observed residual porosity values extracted from new fabrications. It can be seen that after the third iteration, the optimal mean residual porosity is 1.93%, with a corresponding fabrication condition of 163W laser power and 985mm/s scanning speed. The standard deviation showing the uncertainty of the prediction has been reduced to 1.21%, which indicates a higher prediction accuracy.Table 2. The results for Bayesian optimization at each iteration. For each iteration, a predicted optimal mean residual porosity value and its standard deviation are given.

The results for Bayesian optimization at each iteration. For each iteration, a predicted optimal mean residual porosity value and its standard deviation are given. New fabrications are conducted with these predicted optimal process parameters. The porosity values are observed from these newly fabricated parts and used to update the surrogate model in Bayesian optimization for the next iteration.Bayesian optimization iteration1 2 3Predicted optimal process parametersLaser power (W) 150 160 163Scan speed (mm/s) 1070 970 985Energy density (J/mm3) 38.94 45.81 45.96Residual porosityPredicted (Std.) (%) 0.78 (4.39) 1.55 (3.21) 1.93 (1.21)Observed (%)2.43~7.63 2.36~6.18 -Meanwhile, the observed residual porosity values have been reduced after the second iteration compared to the initial dataset, indicating that the quality of the nitinol block has been improved due to better fabrication conditions.

Moreover, the optimal process parameters appear to converge after the second iteration with fewer changes than after the previous iteration. More iterations will be helpful to verify the convergence of process parameters.4 Conclusion and future workIn this work, a nonparametric Bayesian framework incorporating prediction and optimization based on the GP and Bayesian optimization is proposed to discover the optimal laser power and scanning speed of the LB-PBF process for high-density nitinol manufacturing. The GP- surrogate model establishes the initial fabrication, and the Bayesian optimization leads to an efficient acquisition of optimal process parameters to fabricate high-density nitinol samples. The optimal process parameter trends to the convergence after a few iterations of the Bayesian framework.

The optimal process parameter trends to the convergence after a few iterations of the Bayesian framework. The framework can reduce the number of experiments in a data-driven manner and advance the innovations with nitinol in additive manufacturing.The authors’ future research will focus on the three aspects of this work:(1) Additional iterations will be conducted to substantiate the convergence of optimization results and obtain the global optimal parameters (i.e., laser power and scanning speed) for LB-PBF fabricated nitinol samples.293(2) More process parameters will be involved in the proposed Bayesian framework, such as hatch distancing and layer thickness, which also affect the relative density of the LB- PBF fabricated nitinol samples. The proposed Bayesian framework could be expanded to more process parameters with a similar process.

The proposed Bayesian framework could be expanded to more process parameters with a similar process. More importantly, including these process parameters could provide a more comprehensive understanding of the LB-PBF process to improve the reliability of the proposed Bayesian framework.(3) The Batch Bayesian optimization method will be incorporated in the iterative process of the framework. The batch Bayesian optimization could provide a batch of different fabrication conditions within an area that may contain the optimal one. Since the batch production at different fabrication conditions is feasible for the LB-PBF process and more cost-effective than fabricating a single block. Thus, the application of batch Bayesian optimization will greatly enhance process parameter optimization for the additively manufactured nitinol.References[1] W. B. Cross, A. H. Kariotis, and F. J. Stimler, Nitinol characterization study. NASA,[2][3][4][5]Langley Research Center, 1969. D.

e BCC unit cell for latticed SLM parts; mechanical properties as a function of cell size.I. Maskery,1 A.O. Aremu,1 M. Simonelli,1 C. Tuck,1 R.D. Wildman,1 I.A. Ashcroft,1 and R.J.M. Hague1 1Faculty of Engineering, University of Nottingham, Nottingham NG7 2RD, UKThe existing framework describing the mechanical properties of lattices places strong emphasis on one important property, the relative density of the repeating cells. In this work, we explore the eﬀects of cell size, attempting to construct more complete models for the performance of lattices. This was achieved by examining the elastic modulus and ultimate tensile strength of latticed parts with a range of unit cell sizes and ﬁxed density. The parts were produced by selective laser melting (SLM). The examined cell type was body-centred-cubic (BCC), a cell of high relevance for SLM because of its self-supporting structure.

The parts were produced by selective laser melting (SLM). The examined cell type was body-centred-cubic (BCC), a cell of high relevance for SLM because of its self-supporting structure. We obtained power law relationships for the mechanical properties of our latticed specimens as a function of cell size, which are similar in form to the existing laws for the density dependence. These can be used to predict the properties of latticed column structures comprised of BCC cells, and may be easily amended for other situations. In addition, we propose a novel way to analyse the elastic modulus data, which may lead to more general models, applicable to parts of varying size.

In addition, we propose a novel way to analyse the elastic modulus data, which may lead to more general models, applicable to parts of varying size. Lastly, our general methodology may be of use in future studies which explore the other parameters that determine lattice performance; the choice of cell type, the global shape of the lattice structure and the type of stress.Keywords: selective laser melting; lattice; aluminium alloyINTRODUCTIONOne of the most promising capabilities of additive manufacturing (AM) is the production of novel lightweight structures which, of course, are in high demand across sectors such as automotive, medicine and aerospace.

Many innovations in this ﬁeld have come from the use of topology optimisation as a design tool, which is typically used to identify the material layout that maximises speciﬁc mechanical properties[1–3].An alternate and complementary route to better parts, one that is enabled by the manufacturing freedoms of AM, is the replacement of otherwise solid volumes with lattice structures. Such designs have the potential to deliver large reductions in part weight, while also oﬀering high levels of energy absorption under static and dynamic loading[4–8]. However, the choice of lattice unit cell design is large and the mechanical properties of diﬀerent cell types are far from fully understood.

However, the choice of lattice unit cell design is large and the mechanical properties of diﬀerent cell types are far from fully understood. This is a ma- jor obstacle in the development of design methods that utilise lattice structures eﬀectively.Beside the choice of cell type (body-centred-cubic (BCC), face-centred-cubic (FCC), gyroidal, hexahedral, to provide a few examples), at least two other cell properties are signiﬁcant in determining its mechanical performance; its relative density (also known as its volume fraction) and its size. The ﬁrst of these factors has received some prior attention, largely because of the similarity between lattice structures and the more well established foams, while the second has received virtually none.688xyzFIG. 1: The BCC unit cell with a relative density, ρ∗, of 0.36.This work examines the elastic modulus and ultimate tensile strength of latticed parts with BCC unit cells of varying size and ﬁxed density.

The parts were produced in Ti-6Al- 4V alloy using selective laser melting (SLM) and the cell type was chosen because it is known to be self-supporting (it includes no horizontal members) and has been examined experimentally and theoretically elsewhere[5, 6, 9, 10]. The main aim of the investigation is to uncover how the mechanical properties of lattice structures vary according to size of the chosen unit cell. We will attempt to explain the origins of these relationships and add our ﬁndings to the semi-empirical models that currently exist.The Gibson-Ashby modelGibson and Ashby analysed the mechanical properties of metal and polymer foams by considering them as systems of open or closed regular cells, expanding upon other work on honeycomb structures[11, 12].

They used analytical methods based on the beam theory of Timoshenko and Goodier[13], and Roark and Young[14] to relate several properties (the elastic modulus and plastic yield strength, amongst others) of a foam under compression or tension to its relative density, ρ∗. The relationships put forward by Gibson and Ashby were seen to describe the mechanical bahaviour of several foams adequately[11, 15, 16]. The equations have since been applied to latticed structures comprising repeating regular unit cells, an application they were essentially devised for[4, 17, 18].Before presenting Ashby and Gibson’s relationships, some properties of a lattice struc-ture must be deﬁned. We have that,ρ∗ = ρlatt./ρsol.,where ρlatt. and ρsol. are the densities of a lattice structure and a fully-dense solid, respec- tively (both composed of the same material). In this way, ρ∗, can be seen simply as the fraction of a particular volume taken up by the solid material of the lattice.

In this way, ρ∗, can be seen simply as the fraction of a particular volume taken up by the solid material of the lattice. Similarly, we can deﬁne,E∗ = Elatt./Esol., σ∗ U = σU latt./σU sol.,so that E∗ and σ∗ U are the elastic modulus and ultimate tensile strength of a lattice represented as fractions of those of a fully-dense solid of the same material. These will henceforth be referred to as the relative elastic modulus and relative ultimate tensile689(1)(2a) (2b)strength.Gibson and Ashby laid down the straightforward scaling relationship for E∗ as a func-tion of the relative density[11, 12, 19];E∗ = C1ρ∗m,where the exponent m varies depending on the relative contributions of stretching and bending in the cellular deformation process; it is given as m = 2 for open, bending- dominated, cell types, with additional terms being added to equation 3 to account for membrane stresses and gas pressure eﬀects in closed cells.

The prefactor C1,“includes all of the geometric constants of proportionality[11],” and therefore varies signiﬁcantly according to the specimens being examined.The issue of ultimate tensile strength is a little more involved, with a simple relationship Instead, for an open-celled lattice underof the form of equation 3 not forthcoming. compression, Gibson and Ashby provided,σ∗ cr ∝ l c(− 3 mw)ρ∗( 32 − 1 mw),for the crushing strength, σ∗ cr, as a function of cell width, lc, and relative density. mw is the Weibull modulus, a parameter used to quantify the variance in strength amongst It is related to the distribution of ﬂaws in the material and, samples of brittle parts. in this case, it principally dictates the size dependence of the part strength; the larger the Weibull modulus, the smaller the reduction in strength due to increasing part size.

in this case, it principally dictates the size dependence of the part strength; the larger the Weibull modulus, the smaller the reduction in strength due to increasing part size. Physically, this is due to the fact that larger parts are more likely to contain larger pre-existing cracks or pores.For an open-celled lattice under tension, Gibson and Ashby similarly provided,K ∗IC ∝ l c( 1 2 − 3 mw)ρ∗( 32 − 1 mw),to describe the fracture toughness, K ∗ IC, again, as a function of cell width and relative density. In this case, mw = 6 is a limit; the fracture toughness increases with cell size if mw > 6 and decreases with cell size if mw < 6.The cell size dependence for the mechanical propertiesPart of the motivation for this work stems from the recognition that Gibson and introduced in equation 3, must subsume virtually all of the Ashby’s parameter C1, geometrical information about the lattice except for its relative density.

The factors which are thought to determine this parameter are; (i) the choice of unit cell type, of which there are a great many with diﬀerent potential applications and varying suitability for SLM (and AM in general), (ii) the global shape of the lattice structure, (iii) the type of stress, and (iv) the size of the repeating unit cell. This work focuses on the last of these properties, posing the question: with ﬁxed relative density, how does the unit cell size aﬀect the elastic modulus of a lattice? Information in response to this question could be used to develop more eﬀective methods for predicting the moduli of latticed parts, and would therefore be highly valuable as a design tool.690(3)(4)(5)Gauge width (mm)Unit cell width (mm)5.00 1.00 1.

˙66 2.507.00 1.003.505.007.00TABLE I: The range of unit cell widths for the two types of test specimen.With respect to the relative ultimate tensile strength of the lattices, a cell size depen- dence is essentially predicted, though the form it takes has not been explicitly laid out. We will attempt to identify this dependence and relate it to the existing theory for the closely related properties from equations 4 and 5. Once again, any knowledge gleaned regarding the performance of lattices with diﬀerent sized cells, and equivalent density, could be useful in the development of methods to design lattices to meet speciﬁc load conditions.EXPERIMENTAL DETAILSSpecimen productionTensile test specimens with square cross-section were designed according to the ISO standard 6892-1:2009. Specimens with two gauge widths, 5 and 7 mm, were examined.

Specimens with two gauge widths, 5 and 7 mm, were examined. The gauge volumes comprised a repeating BCC unit cell of varying size contained within a ‘net skin’ (determined after production to have thickness 0.243 ± 0.007 mm). The net skin was employed so that an extensometer could be ﬁtted to the exterior of the gauge surfaces, allowing accurate strain measurements. The BCC unit cell, which in our work had a relative density of 0.36, is illustrated in ﬁgure 1.The specimen details, that is, the range of unit cell sizes for each tensile bar, are provided in table I. Figures 2 and 3 are CAD representations of the tensile specimens; the former shows both the 5 and 7 mm gauge versions of the part, including dimensions, while the latter provides a comparison of a latticed gauge region with and without the net skin, revealing the BCC cells beneath.

Several parts with solid gauge volumes were also examined, so that the relative mechanical performance of the latticed specimens could be determined.The specimens were provided by Renishaw Plc. They were manufactured from Ti-6Al- 4V using a Renishaw AM250 SLM machine. The laser power, laser scan speed and hatch spacing were 200 W, 600 mm/s and 150 µm, respectively, and a ‘meander’ scanning strat- egy was taken, whereby the hatch path of each subsequent layer was rotated by 67◦ from the previous one. This strategy serves to reduce any anisotropy, geometric or mechanical, that might occur if there were no hatch rotation. The test specimens underwent a stress relieving heat treatment (600 ◦C for 3 hours under an Ar-rich atmosphere) before removal691r = 32 mmr = 23 mm7 mm40 mmxyz50 mm12 mm8.6 mm67 mm5 mm40 mmFIG. 2: CAD drawings of the 5 mm gauge (right) and 7 mm gauge (left) tensile specimens, including dimensions.FIG.

2: CAD drawings of the 5 mm gauge (right) and 7 mm gauge (left) tensile specimens, including dimensions.FIG. 3: CAD drawing of the latticed region of the 5 mm gauge specimen with 1. ˙66 mm unit cells. The region is shown with (left) and without (right) the net skin that was employed so that a surface-mounted extensometer could be used during testing.from the build plate. Figure 4 shows a selection of the 7 mm gauge width specimens prior to testing, including one which is fully solid and two latticed parts with unit cells of width 3.50 and 7.00 mm.692FIG. 4: Photographs of a selection of 7 mm gauge parts. From left to right; solid gauge section, 3.50 mm unit cell and 7.00 mm unit cell. The inset shows a close-up of the latticed regions, with the net skin and internal spars visible.Tensile testingElastic moduli, Elatt., and ultimate tensile strengths, σU latt., were recorded for the latticed 5 and 7 mm test bars.

All measurements were made using an Instron-5969 universal testing machine. The tensile strain was applied at a rate of 0.01 mm/s. An Instron series 2630 extensometer with 25 mm gauge length was clamped to the center of the latticed section of each test specimen in order to record the strain. Elastic moduli were determined from the regions up to 0.5% strain of the experimental stress-strain curves. Measurements were performed on at least two, and usually three, specimens per design (each combination of cell size and gauge width), with the means and standard errors being used for subsequent analysis.The moduli of completely solid (non-latticed) test specimens were also recorded. These were 102.3±0.9 and 101±1 GPa for the 5 and 7 mm bars, respectively, and provide the Esol. of equation 2a. They agree with each other within experimental error and are slightly lower than values typically associated with cast and forged alloys (∼ 105 − 116 GPa[20, 21]).

of equation 2a. They agree with each other within experimental error and are slightly lower than values typically associated with cast and forged alloys (∼ 105 − 116 GPa[20, 21]). The ultimate tensile strength of the solid bars was found to be 1.07 ± 0.01 GPa, slightly higher than expected from cast alloy but lower than has been reported for some Ti-6Al-4V SLM parts[21–23]. This value constitutes σU sol. of equation 2b.RESULTS AND DISCUSSIONNumerical results from the mechanical testing of all specimens are provided in table U , are as deﬁned in equations 2a and 2b. TheII, where the relative properties, E∗ and σ∗693lc (mm)vc (mm3)v∗ ×10−3Elatt. (GPa)E∗ ×10−3σU latt. (MPa)σ∗ U ×10−35 mm gauge bars1.00 1.

The relative properties (v∗, E∗, and σ∗U ) are deﬁned in the text.unit cell volumes, vc, for each specimen type are also included; for BCC cells these are simply the unit cell width cubed. v∗ is a property which will be deﬁned and discussed shortly.The elastic modulusThe relative elastic moduli are plotted in ﬁgure 5 as a function of cell width. The E∗ values decrease over the examined range of cell width, from (19.4 ± 0.5)% to (7.0 ± 0.5)% of the Esol. values obtained from solid bars. The second feature of note is that the data from both the 5 and 7 mm test bars follow the same distribution, signifying that we may treat them as a single data set for the purposes of data analysis.We ﬁrst hypothesize that E∗ follows a power law with lc similar to that already knownto exist for ρ∗. We suggest,E∗ = U1 l n1c,where U1 and n1 are analagous to Gibson and Ashby’s C1 and m of equation 3. The resulting weighted least-squares ﬁt is shown as a dashed line in ﬁgure 5.

We suggest,E∗ = U1 l n1c,where U1 and n1 are analagous to Gibson and Ashby’s C1 and m of equation 3. The resulting weighted least-squares ﬁt is shown as a dashed line in ﬁgure 5. The values of U1 and n1 were found to be 0.179 ± 0.009 and −0.56 ± 0.05, respectively. However, this ﬁt is statistically rather poor, with a ¯χ2 (reduced-χ2) value of 2.29, and the ﬁt residual, shown in the upper panel of ﬁgure 6, shows signiﬁcant additional structure. Clearly, this simple model is incapable of describing the data set well.Two variations on this scaling law were also trialled; one with a constant to modify the cell width, the other with a constant to modify the relative elastic modulus. Physically, such constants could be interpreted as systematic errors associated with the experimental procedure, or they might be true requirements of the model required to describe the data.

Physically, such constants could be interpreted as systematic errors associated with the experimental procedure, or they might be true requirements of the model required to describe the data. The modiﬁed power laws are given by E∗ = U1 (lc + δlc)n1 and E∗ = U1 l n1 c + δE∗, and are shown in ﬁgure 5 as solid and dotted lines, respectively. As anticipated, the ¯χ2 values associated with these ﬁts are much improved, with the former providing a value very close to unity. Furthermore, the central panel of ﬁgure 6 shows far less residual structure for694(6)0.2020.06450.14lc(mm)E∗0.1817mmgaugebars70.10E∗=U1ln1c+δE∗¯χ2=1.18E∗=U1(lc+δlc)n1¯χ2=0.9830.12E∗=U1ln1c¯χ2=2.290.080.165mmgaugebars6FIG. 5: The relative elastic modulus, E∗, of the 5 and 7 mm latticed test bars as a function of cell width, lc. Weighted least-squares power law ﬁts are shown as dashed, solid and dotted lines.

5: The relative elastic modulus, E∗, of the 5 and 7 mm latticed test bars as a function of cell width, lc. Weighted least-squares power law ﬁts are shown as dashed, solid and dotted lines. The ﬁtting procedure is described in the text.the E∗ = U1 (lc + δlc)n1 ﬁt, compared to the the unmodiﬁed form. The parameters from this ﬁt were U1 = 0.11±0.01, δlc = −0.8±0.1 mm and n1 = −0.28±0.07. The determined cell width oﬀset of -0.8 mm could be indicative of a systematic error in the manufacture of the latticed bars. However, this issue was subsequently investigated; the cell widths were found to conform well to the expected values.The second modiﬁed power law, E∗ = U1 l n1c + δE∗, provides a poorer ﬁt than that described above, with ¯χ2 = 1.18 and more residual structure, but it is still superior to the unmodiﬁed form.

The ﬁtted parameters are U1 = 0.12 ± 0.01, n1 = −1.3 ± 0.4 and δE∗ = 0.06 ± 0.01.Anomalous behaviour of the smallest cellsFurthering the analysis of the E∗(lc) data, we speculate that data from 1 mm cells in both the 5 and 7 mm latticed test bars may be outliers or simply not described by the same mathematical model that describes the rest of the data set. Evidence for this assertion comes from the residuals in ﬁgure 6, where the 1 mm data are almost equally poorly described by all three models trialled so far. There may be a physical cause for this; the cavities, or voids, in the 1 mm lattice structures are small enough for the complete removal of Ti-6Al-4V powder to be problematic. If residual powder remained6951E∗=U1(lc+δlc)n165mmgaugebarslc(mm)E∗residual0.01574−0.015E∗=U1ln1c+δE∗−0.015057mmgaugebars−0.01520.015030.015E∗residual0E∗=U1ln1cE∗residualFIG.

6: Residuals from the relative elastic modulus ﬁts of ﬁgure 5.in some or many of the internal lattice cavities for the 1 mm cells, the behaviour of those bars under tension may be quite diﬀerent from the others. Another possible explanation is that the performance of these cells is being aﬀected by the presence of loosely bound powder around the lattice struts and nodes. Loosely bound, or partially sintered, powder is a persistent issue in SLM and aﬀects diﬀerent materials to diﬀerent extents, often its presence is related negatively impacting properties such as the surface roughness; to the choice of laser processing parameters and the morphology of the powder. In the case of our latticed structures, it would have a larger eﬀect for smaller cells, where the dimensions of the struts and nodes are smallest.The relative elastic moduli were re-ﬁt with E∗ = U1 l n1, this time excluding those data from the 1 mm cells; the result is shown in ﬁgure 7.

The parameters are U1 = 0.144±0.002 and n2 = −0.130 ± 0.004. The ¯χ2 value is extremely low at 0.02, most likely indicating that, with the exclusion of two data points (and the reduction to just three degrees of freedom for the ﬁtting procedure), the ﬁt is over-paramaterised.cAlternate analysisIn ﬁgure 8 we present E∗ as a function of the relative cell volume, v∗. This, we deﬁneas,v∗ = vc/vlatt. env.,696(7)0.1265mmgaugebars3507mmgaugebars70.160.015E∗=U1ln1c(1mmdataexcl.)¯χ2=0.02−0.015lc(mm)E∗residual0.080.1440.200.0610.10E∗20.18FIG. 7: 5 and 7 mm E∗ data ﬁt with E∗ = U1 l n1 the ﬁt, as described in the text. The lower panel shows the corresponding ﬁt residual.. Data from the 1 mm cells are excluded fromcwhere vc is the cell volume (given simply by l 3 c ) and vlatt. env. is the volume of the latticed environment in which the cell resides. For the 5 and 7 mm bars, vlatt. env.

env. is the volume of the latticed environment in which the cell resides. For the 5 and 7 mm bars, vlatt. env. has values 875 and 2401 mm3, respectively (i.e. 5 × 5 × 35 mm and 7 × 7 × 49 mm, where 35 and 49 mm are the lengths of the latticed gauge sections of the bars).Choosing to work with a relative measure of cell size like v∗, rather than the absolute values of lc or vc, may provide a way to formulate general descriptions of lattices that are independent of the size of the part, equally applicable to parts of quite diﬀerent length scales. There is precedent for this methodology in the way we work with E∗, for example, instead of E. This is done so that general rules, such as equations 3, 4 and 5, may be developed that apply to a range of materials, including polymers, ceramics and metals, where the absolute moduli diﬀer greatly.

Similarly, ρ∗ is used in- stead of ρ because the absolute densities of the materials in question also vary signiﬁcantly.An unmodiﬁed power law was applied and it is shown as a solid line in ﬁgure 8. Weighted least-squares ﬁtting provided the result; E∗ = (0.051 ± 0.004) v∗(−0.17±0.01) with ¯χ2 = 1.68. The ﬁt is statistically poorer than the previous best, but it has the advantage of describing the E∗ behaviour equally well over the full range of cell size, including the smallest cells.6970.060.120.120.04v∗¯E∗0.080.080.140.1020.0085mmgaugebarsweightedleast-squaresﬁt6100.080.040.1612x 10−30.0040.16E∗∝v∗(−0.17±0.01)¯χ2=1.68¯E∗∝v∗(0.89±0.06)07mmgaugebars0.12v∗E∗FIG. 8: The relative elastic modulus, E∗, of the 5 and 7 mm latticed test bars as a function of relative cell volume, v∗.

8: The relative elastic modulus, E∗, of the 5 and 7 mm latticed test bars as a function of relative cell volume, v∗. The inset plot shows the relative elastic modulus per unit cell, ¯E∗, also as a function of v∗.The ultimate tensile strengthThe relative ultimate tensile strengths, σ∗U , of the 5 and 7 mm latticed bars are shown in ﬁgure 9. They take values similar to those seen for E∗, that is, decreasing from (17.7 ± 0.2)% to (5.2 ± 0.2)% of the strengths of fully-dense test bars. Following the same approach as outlined above, three power laws, two with additional constants for lc and σ∗ U , were trialled; they are shown as dashed, solid and dotted lines.U = U2 (lc + εlc)n2 that provides the best ﬁt, with the unmodiﬁed power law performing signiﬁcantly worse. The factors are U2 = 0.7±0.6, εlc = 2±1 mm and n2 = −1.2±0.3.

The factors are U2 = 0.7±0.6, εlc = 2±1 mm and n2 = −1.2±0.3. The ¯χ2 values for the three trialled ﬁts are larger than obtained in the E∗ analysis, reﬂecting statistically poorer ﬁts overall; this is due mainly to the smaller fractional errors associated with the σU measurements. This has also given rise to larger uncertainties on the determined parameters U2, εlc and n2. Interestingly, switching the dependent variable to the relative cell volume, v∗, instead of lc, as implemented previously, did not provide an improved ﬁt when using the unmodiﬁed power law, σ∗Once again, it is the power law of the form σ∗U ∝ v∗n.6986σ∗U=U2(lc+εlc)n2¯χ2=3.1970.06lc(mm)σ∗U0.040.08340.100.1210.187mmgaugebarsσ∗U=U2ln2c+εσ∗U¯χ2=3.472σ∗U=U2ln2c¯χ2=8.4250.140.165mmgaugebarsFIG. 9: The relative ultimate tensile strength, σ∗ U , of the 5 and 7 mm latticed test bars as a function of cell width, lc.

9: The relative ultimate tensile strength, σ∗ U , of the 5 and 7 mm latticed test bars as a function of cell width, lc. Again, a weighted least-squares power law ﬁt, shown as a solid line, provides the dependence of σ∗U on lc.The Weibull modulusWe can use the determined parameter n2 from our best ﬁt to estimate the Weibull modulus of our latticed parts. We speculate that the cell width dependence for σ∗ U is likely to be, or be similar to, one of those seen in equations 4 and 5. Thus, our σ∗ U values would decrease with cell size according to l . The ﬁrst case yields c mw = 2.5 ± 0.6, whilst the second provides mw = 1.8 ± 0.4. Reasonable sources of comparison for these values are understandably scarce in the literature, owing to the lack of research on lattice structures produced by SLM. The most meaningful results currently available may be the work of Blazy et al.

The most meaningful results currently available may be the work of Blazy et al. and Ramamurty et al., who provide values of mw between 8 and 16.7 ± 0.6 for some aluminium foams[24, 25], but since both the material and the manufacturing process in those works are diﬀerent from our own, direct comparison is not advised. Clearly, understanding how our values of mw relate to other work is worthy of further research.− 3 mw( 1 2 − 3 mw)or l cCONCLUSIONSThe main novel ﬁnding of this investigation is that there is a dependence of the relative elastic modulus of lattice structures on the size of the cells they comprise. Our best ﬁt suggests E∗ ∝ l (−0.28±0.07) , though the extent to which this proportionality is applicable to lattices with diﬀerent unit cell types and diﬀerent global geometries is notc699yet fully understood.

This ﬁnding is signiﬁcant because previous investigations into lattices have chieﬂy focussed on the inﬂuence of relative density, with any cell size eﬀect forming part of a constant, C1, to be determined for each individual case. Our result could be used to develop more complete predictive models for the design of latticed parts; at its simplest it suggests a clear advantage in choosing smaller lattice unit cells if maximum stiﬀness is required. However, there are two important points of note; (i) a constant, which we call δlc, was required to modify a basic power law to provide the best ﬁt to our data, and (ii) even our modiﬁed power laws failed to properly describe the data from the test specimens with the smallest cells. The smallest cells may be outliers, or they may actually behave diﬀerently to the others.We have shown that a possible solution to these issues lies in analysing the E∗ data as a function, not of the cell width, but of the relative cell volume, v∗.

This methodology provided a reasonable ﬁt with a simpler power law than was previously achievable. An advantage of this analytical method is that it may provide a way to generalise the perfor- mance of lattices of diﬀerent sizes, in much the same way that the well established use of E∗ and ρ∗ allows us to directly compare lattices of diﬀerent materials. If this relationship proves similarly successful in further investigations, it may ultimately be found that the relative elastic modulus of lattices can be described generally by,E∗ = Um ρ∗m1v∗m2,where Um is dependent only on the choice of unit cell and the global shape of the part. The implication would be that a latticed cube composed of 1000 unit cells would have the same relative elastic modulus regardless of its size, be it 20 mm wide or 100 mm wide.

The implication would be that a latticed cube composed of 1000 unit cells would have the same relative elastic modulus regardless of its size, be it 20 mm wide or 100 mm wide. This requires further investigation.Lastly, we found a cell size dependence for the ultimate tensile strength of our latticed bars; the data followed σ∗ , though, once again, the best ﬁt was achieved only with an lc oﬀset, εlc. The choice of lc as the independent variable was appropriate for the analysis of σ∗ U data; there were no signiﬁcant outliers at low cell width, and there was no improvement of ﬁt when v∗ was used instead. As above, the major implication of this ﬁnding is that smaller lattice cells should be chosen for applications where high tensile strength is required. We used our determined lc exponent to estimate the Weibull modulus of our lattices, making the assumption that σ∗ U follows a law similar to those laid out by Gibson and Ashby for crushing strength and fracture toughness.

We found mw = 2.5 ± 0.6 or mw = 1.8 ± 0.4, which are relatively small, but the lack of available sources for comparison makes a proper assessment impossible.U ∝ l (−1.2±0.3)CAcknowledgmentsWe would like to thank Sean Smith, Ravi Aswathanarayanaswamy, Hossein Sheykh- Poor and various teams at Renishaw Plc. for the provision of the test specimens and continued support in our collaborative projects. Thanks also to Mark East, Mark Hardy and Joe White, technicians of the Additive Manufacture and 3D Printing Research Group at Nottingham. Funding was provided by the UK Technology Strategy Board (TSB).700(8)[1] G. Chahine et al., in Solid Freeform Fabrication Symposium, 2010.[2] D. Brackett et al., in Solid Freeform Fabrication Symposium, 2011.[3] N. Gardan, Int. J. Manuf. Eng. 2014 (2014).[4] C. Yan et al., Mater. Design 55, 533 (2014).[5] M. Smith et al., Int. J. Mech. Sci. 67, 28 (2013).[6] K. Ushijima et al., J. Sandw. Struct. Mat. 13, 303 (2011).[7] N. A.

am Deletion in Square Honeycombs for Improved Energy Absorption Under Quasi-Static In-Plane CompressionIrving E. Ramirez-Chavez*, Christine Lee†, Dhruv Bhate*3DX Research Group, Arizona State University, Mesa, AZ 85212 †Interwoven Labs, Tempe, AZ 85282AbstractWhen selecting cellular materials for energy absorption applications, there have traditionally been two choices: a periodic structure such as a honeycomb, or a stochastic one, as seen in foams. Both choices involve a global definition governing the allocation of the members of the structure, be they beams or surfaces. With Additive Manufacturing, the exploration of more complex structures enables the creation of aperiodicity through the local modification of periodic structures. This paper explores one application of this approach by deleting beams in square honeycombs, with the aim of avoiding localization of failure that generates significant undulations in the stress plateau under in-plane quasi-static compression.

These perturbed structures show improved energy absorption behavior by generating higher Specific Energy Absorption for a given transmitted stress and relative density than their periodic counterparts. This work thus argues for further exploration of localized aperiodicity as an approach to finely tune energy absorption performance.Keywords: Additive Manufacturing, beam deletion, cellular materials, densification strain, maximum transmitted stress, perturbation, energy absorption.IntroductionNature has an intriguing way of leveraging cellular materials in its many forms, many of which have served as inspiration for engineering design [1]. Structures like bone and insect nests integrate a solid and an empty phase to generate multifunctionality, which have been explored in different fields. One such field is art, in which notions of symmetry, negative space, and periodicity have been studied in the context of form and design [2].

One such field is art, in which notions of symmetry, negative space, and periodicity have been studied in the context of form and design [2]. In this field, not only is the distribution of the solid area emphasized, but also how it blends and interacts with the empty region, also known as negative space. In engineering, this interaction of solid and negative phases has a myriad of applications, one of which is energy absorption. Some examples in nature where negative space is optimized is in the peel of the pummelo. Pummelo is a fruit that can grow on trees at a height of 25 feet above ground [3]. When the fruit has ripened, it falls but is not damaged by the impact. A closer examination of the peel of this fruit shows a graded and stochastic cellular pattern, which has been related to its impact damping properties [4]. This phenomenon can be observed in animal organs as well, such as in a Toucan beak or a Woodpecker skull [5], [6].

This phenomenon can be observed in animal organs as well, such as in a Toucan beak or a Woodpecker skull [5], [6]. Both structures have an2106aperiodic open-cell structure similar to that seen in spongy bone [7]. All of these structures have been associated with enhanced energy absorption.Historically, honeycombs and foams have been the primary cellular materials that have found use in energy absorption applications [8]. With the advent of Additive Manufacturing (AM) technologies, research into energy-absorbing cellular materials is exploring different design concepts to enhance their properties. There are several different metrics that characterize an energy-absorbing cellular material [9][10]. From a designer’s perspective, an energy absorber is a material that should be able to stay beneath a particular stress threshold and maintain it until material densification, while maximizing the energy absorbed in the process [11].

An ideal energy absorber should thus satisfy three main characteristics:1. Constant crushing stress 2. High densification strain 3. High Specific Energy Absorption (SEA) per unit mass for a given Maximum Transmitted Stress (MTS)Different cellular design strategies have been explored to approach ideal behavior. For example, the use of a pre-crushed Aluminum honeycomb can substantially decrease the MTS while maximizing SEA due to a more stable plateau stress [12]. The use of metallic and polymer foams has also been explored since stochastic materials are associated with good energy absorption properties in the context of uncertainty in loading direction, similar to spongy bone found in nature. Other examples of design innovation for energy absorption can be found in studies with hollow twin hemispheres [13], which allowed for higher densification strains with low compression modulus.

Other examples of design innovation for energy absorption can be found in studies with hollow twin hemispheres [13], which allowed for higher densification strains with low compression modulus. Other designs include the use of lattices that can spread impact loading over time and reduce the first maximum stress [14], and micro-lattices that enable tunability with high energy absorption and longer densification with a tradeoff of poor constant crushing behavior [15], [16].This study builds on prior work [17] where the idea of breaking symmetry was explored in square grid honeycombs. Symmetry has been explored in the field of mathematics, nature, and art, and is described by mathematicians to be a measure of “invariance to any of various transformations such as rotation, reflection or scaling” [18], [19]. In the prior study, aperiodic honeycombs were designed arbitrarily, as shown in Figure 1, to study what effect they have on large deformation behavior under in-plane compression.

In the prior study, aperiodic honeycombs were designed arbitrarily, as shown in Figure 1, to study what effect they have on large deformation behavior under in-plane compression. For the purpose of the current study the main idea was to design aperiodicity in a more deterministic manner by deleting beams in a 2D square grid honeycomb following a triangular waveform approach and examine how it can influence the mechanical properties. The following sections discuss the approach, the results and conclusions associated with the current study, while also drawing comparisons to prior work.2107Figure 1. Design strategies using beam deletion, reproduced from prior work [17]Design ApproachI. Beam Deletion ApproachA classification of aperiodic architected cellular materials was previously proposed, wherein three main types of aperiodicity were identified: gradation, perturbation, and hybridization [20]. Hybridization was shown to have two main characteristics: a feature and a method.

Hybridization was shown to have two main characteristics: a feature and a method. For this study, the beam member was selected as the feature and the method used to achieve hybridization was deletion. By following this approach, a square honeycomb’s appearance would start to change and generate a hybrid cellular structure, an example of which is shown in Figure 2.Figure 2. Beam deletion technique for generating aperiodicity, reproduced from [20]2108Prior work developed two key insights. The first of these emerged after comparing the compression stress-strain response of baseline regular square honeycomb against a form with an embedded spiral with 73 deleted beams, as shown in Figure 3. Beam deletion has the effect of reducing relative density, which tends to be associated with higher densification strain. However, the connected spiral nature of the structure also resulted in lower First Peak Stress (FPS) and a lower MTS, with a reduced undulation in the plateau region.

However, the connected spiral nature of the structure also resulted in lower First Peak Stress (FPS) and a lower MTS, with a reduced undulation in the plateau region. These are all promising directions towards an ideal energy absorber.Figure 3. Spiral channel dropping peak stress and prolonging densification, adapted from [17]The second important insight and one that inspired the current work, is demonstrated in Figure 4. The red lines connect centers of cells with deleted beams. The deletion of beams in a particular row was found to initiate early failure and therefore a lower FPS, but having too many empty cells in a row resulted in lower energy absorption because the stress levels remained low. The structure with the highest SEA at the lowest MTS was found to be the one in the top left corner of Figure 4, with only two rows connecting empty cells.

The structure with the highest SEA at the lowest MTS was found to be the one in the top left corner of Figure 4, with only two rows connecting empty cells. The best performing structure with regard to maximizing SEA for a given MTS, is reproduced in Figure 5, in comparison to the baseline. Failure bands were observed in both specimens, but the second specimen with staggered empty cells, had bands that formed at an angle instead of the horizontal row by row collapse of the baseline square grid, with each collapse corresponding to a load drop. Thus, a hypothesis was formulated that staggering beam deletion, in such a manner as to avoid collinear empty cells, is likely to improve energy absorption. This is the hypothesis that the current work examines.2109Figure 4. The hypothesis of the load transfer mechanism, reproduced from [17]Figure 5.

This is the hypothesis that the current work examines.2109Figure 4. The hypothesis of the load transfer mechanism, reproduced from [17]Figure 5. Asymmetric member deletion suggests promising behavior, reproduced from [17]2110II.Triangular Wave RationaleFollowing the observations made from the previous study, a design strategy was developed to enhance the energy absorption properties of the square honeycomb by exploiting the observation of staggered negative space and avoiding collinear deletion of beams. To minimize edge effects, two columns of square cells were not modified in any specimen on either side of the specimen, as shown in Figure 6. Since prior results showed deleting four beams at a time showed promising results, this was the strategy employed here as well. Finally, vertical (V) and horizontal (U) spacing between the centers of empty cells was defined as shown in Figure 6, for the purpose of placing the negative spaces in a systematic way.

Finally, vertical (V) and horizontal (U) spacing between the centers of empty cells was defined as shown in Figure 6, for the purpose of placing the negative spaces in a systematic way. U and V thus enable the formation of triangular wave- like pattern, which it was hypothesized would guide the failure band formation without having large reductions in stress during the compression event, and a more stable plateau.Figure 6. Triangular wave beam deletion designFigure 7 shows all the triangular wave designs developed for this study, with their respective U and V values. A total of 16 triangular wave designs and 2 baselines were designed. It can be observed that the triangular wave is more apparent in some specimens than others because of the spacing values and bounding box size. The bounding box sizes are different in some cases which was done to ensure completeness of the design and avoid empty cells of varying sizes.

The bounding box sizes are different in some cases which was done to ensure completeness of the design and avoid empty cells of varying sizes. This discrepancy was addressed in the data analysis, by the use of stress and strain normalization based on the specimen width and depth (to compute area for stress calculation) and specimen height (for strain). In all cases, at least 12 cells were defined in all directions, to minimize size effects, per ISO 13314.2111Figure 7.

In all cases, at least 12 cells were defined in all directions, to minimize size effects, per ISO 13314.2111Figure 7. Triangular wave designs used in this work, with different U and V spacingMethodsThe methodology followed in this study is summarized in Figure 8 and involved the following steps: (i) CAD designs were created with deleted beams and exported to a Stratasys FortusTM 450mc printer which manufactured the specimens using fused deposition modeling (FDM) process with ABS (Acrylonitrile Butadiene Styrene); (ii) This was followed by mechanical testing on an Instron 8801 using a compression setup and an effective strain rate of 10-3 s-1; and (iii) analysis of the data was then performed to extract metrics of interest. The following sections detail these steps.I. Design & ManufacturingA total of 16 designs and 2 baselines were designed as shown in Figure 7 using the aforementioned triangular wave design approach.

Design & ManufacturingA total of 16 designs and 2 baselines were designed as shown in Figure 7 using the aforementioned triangular wave design approach. CAD designs were generated using SolidWorksTM software and then imported to a Stratasys FortusTM 450mc 3D printer. ABS was the selected material, and the single contour printing parameters were used. The specimens were designed with a constant 1-inch thickness so as to have enough stability against buckling during in-plane compression.2112Figure 8. Three main steps used in the study: design and manufacturing, mechanical testing, and computation of energy absorption metrics.II. Mechanical TestingThe honeycombs were tested under quasistatic compression with an effective strain rate of 10-3s-1 on an Instron 8801 machine with a 50kN load cell, following ISO 13314. Tests were conducted till densification with a load-based test stop criteria.

Tests were conducted till densification with a load-based test stop criteria. This test results in force/displacement data that can be used to compute energy absorption metrics for analysis. Since the specimens had slightly different sizes, this force/displacement data was converted to stress/strain, in order to, be able to normalize and compare properties among them.III. Computation of Energy Absorption MetricsOnce the stress/strain data is obtained, it was plotted to obtain graphs similar to that shownin Figure 9. MATLAB was then used to obtain the following two metrics:a) First Peak Stress and Maximum Transmitted StressThe Maximum Transmitted Stress (MTS) is the highest stress reached prior to densification, indicated in Figure 9 with the green marker, and is typically normalized by the material yield stress.

MTS may or may not be the first peak (FPS), indicated with the red marker, and is of greater interest from a designer’s perspective.b) Specific Energy Absorption (SEA)This is defined as the area under the curve from the origin to strain densification (ɛD) and then normalized by the specimen density (ρ). It is shown as the shaded blue region in Figure 9 and estimated as:𝜀𝜀𝐷𝐷 ∫ 𝜎𝜎𝜎𝜎𝜎𝜎 0(1)𝑆𝑆𝑆𝑆𝑆𝑆 =𝜌𝜌2113Figure 9. Mechanical properties of interest for energy absorptionResultsI.Comparisons between prior and current workA commonly used approach to identify promising cellular material candidates for energy absorption is to plot SEA vs FPS, with the preferred options having high SEA at low FPS values. The data from the prior study from 2018 [17] was combined with the results from this study and are shown together in Figure 10.

The data from the prior study from 2018 [17] was combined with the results from this study and are shown together in Figure 10. Baseline square and spiral specimens were tested in the current study to ensure no significant deviation in print quality or testing conditions and are shown encircled at the extremes of the data collected. The majority of the results, including all results from the prior study, fell on a line connecting these extremes, but crucially from this study, it was found that most of the triangular wave designs were remarkably clustered together at a higher level of SEA relative to expectation for the FPS associated with them. This suggests that the triangular wave design approach is a favorable allocation of negative space for improving energy absorber design.Figure 10. SEA vs FPS comparison between this work and prior [17]2114II.

SEA vs FPS comparison between this work and prior [17]2114II. SEA and MTSAn examination of the SEA and MTS values for the specimens in the current work, relative to their U and V values is shown in Figures 11a and 11b, respectively. The asterisk marker represents the baseline periodic square grid honeycomb. Several designs have higher SEA and lower MTS than the baseline, but in particular, a few points have higher SEA than the trend, and are worthy of further examination.(a)(b)Figure 11. a) SEA vs relative density, and (b) MTS vs relative density for the triangular wave designs studied in this work, the asterisk (*) represents the periodic square baselineIn Figure 12, a more meaningful, designer’s plot which shows SEA vs MTS, allows the identification of high performing structures with high SEA for a given MTS (top left corner).

When observing Figures 11 and 12 together, it is remarkable how values of horizontal spacing U=2 (represented by the light blue markers), tends to have the best performance across different V values.Figure 12. SEA vs MTS for the triangular wave designs studied in this work, the asterisk (*) represents the periodic square baseline2115To try and understand why the U=2 specimens had higher SEA values, it is useful to study the nature of their compression response relative to the baseline, as shown in Figure 13. From the stress/strain plots it can be seen how undulation is reduced and densification strain is extended, while conserving a lower FPS and MTS, with the net effect of improved SEA.Figure 13.

From the stress/strain plots it can be seen how undulation is reduced and densification strain is extended, while conserving a lower FPS and MTS, with the net effect of improved SEA.Figure 13. U2V2 specimen compression response compared to the periodic square baselineIII.Significance of Statistical Correlations The prior discussion demonstrated that the introduction of triangular wave approach to negative space improves SEA by greatly reducing the FPS and MTS, pushing out densification strain, and reducing undulations. To assess whether the U and V factors do indeed correlate to observations, a bivariate fit was done and p-value assessed for the significance of the fit, for SEA and SEA/MTS in Figures 14 and 15, respectively. The latter ratio was chosen to condense the designer’s perspective into a single metric, where higher values of SEA/MTS are desirable. A caveat with this analysis is that it is conducted with one variable (U or V) at a time only.

A caveat with this analysis is that it is conducted with one variable (U or V) at a time only. Given this limitation, any significance is particularly remarkable – and it was found to be significant (p- value less than 0.05 in only one instance, viz. for SEA/MTS versus U, in Figure 15. This adds more weight to the prior observation that the horizontal spacing between empty cells is a more important variable than the vertical spacing, in improving energy absorber performance, for the triangular wave design approach proposed here.2116Figure 14. Bivariate analysis for SEA as a function of U and V separatelyFigure 15.

Bivariate analysis for SEA as a function of U and V separatelyFigure 15. Bivariate analysis for SEA/MTS as a function of U and V separately2117ConclusionsThe current work proposed a new triangular wave design derived from beam deletion to create negative space that resulted in a cluster of designs in which the majority showed improved energy absorption performance more than baseline regular square honeycombs, and prior work with arbitrarily distributed negative space. This was determined by a study of the effect of these beam deletion designs on Specific Energy Absorption, stresses (MTS and FPS) and densification strain. The triangular wave approach of beam deletion influences the failure band trajectory which directly impacts the overall plateau stress undulation, densification, and in some cases, reduces the MTS, and FPS, and increases the SEA/MTS.

With regards to the latter parameter, statistical analysis suggests the key design parameter that influences SEA/MTS is the horizontal spacing between beam deleted empty cells. More work is needed to further explore the role of deleting cellular members in influencing mechanical behavior, as one of many strategies to explore aperiodicity in these fascinating materials.References[1] Y. Mistry, D. Anderson, and D. Bhate, “Bio-Inspired Design,” in Fundamentals of Laser Powder Bed Fusion of Metals, 1st ed., E. M. Igor Yadroitsev, Ina Yadroitsava, Anton Du Plessis, Ed. Elsevier, 2021.[2] W. Wong, Principles of Form and Design, 1st ed. Wiley, 1993.[3]S. F. Fischer et al., “Pummelos as concept generators for biomimetically inspired low weight doi: structures with 10.1002/adem.201080065.excellentdampingproperties,”2010.[4] A.

am Structure Optimization for Additive Manufacturing basedon Principal Stress LinesYongqiang Li, Yong Chen* Epstein Department of Industrial and Systems Engineering University of Southern California, Los Angeles, CA 90089Corresponding author: yongchen@usc.edu, (213) 740-7829ABSTRACTThe benefits of component design with cellular structures have been demonstrated in a wide variety of applications. The recent advances in additive manufacturing and high performance computing have enabled us to design a product component with adaptive cellular structures to achieve significantly better performance. However, designing a product component with such structures, especially its shape and topology, poses significant challenges. Many approaches in topology optimization have been developed before for the purpose. In this paper, we present a novel structural optimization method based on the principal stress line analysis of a continuum domain.

In this paper, we present a novel structural optimization method based on the principal stress line analysis of a continuum domain. We first present the theoretical basis of our optimization method. We then discuss the properties of principal stress lines and their computation in a given design domain. Accordingly a novel structural optimization method is presented including size, shape and topology optimization. Related mathematical formulations and algorithms are also given for generating a beam structure with the minimum compliance. Three test cases are presented to illustrate the presented method.KEYWORDS: Topology optimization, principal stress line, beam structures, minimum compliance.1. INTRODUCTIONOver the last twenty years, solid freeform fabrication (SFF) has been developed based on the layer-based additive principle.

INTRODUCTIONOver the last twenty years, solid freeform fabrication (SFF) has been developed based on the layer-based additive principle. Recent advances in material, process and machine development have enabled the SFF processes to evolve from prototype usage (rapid prototyping) to direct product manufacturing. Examples such as Boeing’s F/A-18 components and Siemens’s hearing aid shells have been reported for applications in aerospace and medical industries. SFF is a direct manufacturing process that can fabricate parts directly from CAD models without part-specific tooling or fixtures. Therefore, for the first time in history, we have a set of manufacturing processes that can cost-effectively fabricate truly complex 3-dimensional shapes.

Therefore, for the first time in history, we have a set of manufacturing processes that can cost-effectively fabricate truly complex 3-dimensional shapes. As identified in (Bourell, et al., 2009; Hopkinson, et al., 2006), a primary advantage of SFF is its capability to allow revolutionary new designs when complex geometry is no longer a limiting factor.Even though SFF provides tremendous design freedom that was unavailable before, the capability of using SFF’s unlimited geometric capabilities for better design is still limited and mainly untapped. As pointed out in (Bourell, et al., 2009), it is crucial to develop new design methods and related CAD tools in order to fully utilize the design freedom provided by SFF.

As pointed out in (Bourell, et al., 2009), it is crucial to develop new design methods and related CAD tools in order to fully utilize the design freedom provided by SFF. Since beam structure is a type of design that can well demonstrate the SFF’s geometric capability, we focus on the design optimization of beam structures in this paper, including those of size, shape/geometry and topology.Topology optimization of trusses is a classical subject in structural design. The study of fundamental properties of optimal grid-like continua was made by Michell (1904). In the last three decades, a considerable amount of work has been carried out on structural optimization. Two developed computational methods that are the most well-known for truss topology design are the Homogenization and the Ground Structure methods. More detailed description of them and related methods are given in Section 2.

More detailed description of them and related methods are given in Section 2. Even though the two methods are well-established, problems such as the generated design needs to be interpreted and in many cases unnatural to users have also been noticed. In addition, both methods are computationally expensive; hence it is rather difficult for designers to control the generatedresults. Motivated by Michell’s work, we present an intuitive method of designing beam structures with a minimum compliance or maximum stiffness. Our method is based on the principal stress line analysis of the given design domain. The mathematical foundation of our method is presented in Section 3. A high- level comparison of our approach with the homogenization and ground structure methods is given in Figure 1. We believe our approach is much simpler and faster.

A high- level comparison of our approach with the homogenization and ground structure methods is given in Figure 1. We believe our approach is much simpler and faster. In addition, the generated structures have no issues that are common to the homogenization and ground structure methods such as gray area and disconnected structures as pointed out by (Lu and Kota, 2006).Figure 1. A comparison of our method with Homogenization and Ground Structure methods (Lu and Kota, 2006).2. RELATED WORKTopology optimization is a classical subject in structural design. Introductions to truss topology problems can be found in (Topping, 1993; Kirsch, 1993; Rozvany, et al., 1995a; Achtziger, 1997; Bendsøe and Sigmund, 2002). Topology optimization of discretized and continuum structures are two broad categories in the structural optimization. We focus on discrete structure optimization in this paper.

Topology optimization of discretized and continuum structures are two broad categories in the structural optimization. We focus on discrete structure optimization in this paper. Hemp and Chan (1970), and Dorn, Gomory and Greenberg (1964) considered a ground structure to overcome the infeasibilities of Michell structures. Given the same design domain, the analogous boundary conditions and external loads, they obtained the trusses coincide with the principal stresses directions of an optimal continuum structure (Achtziger, 1997). Nowadays the ground structure method is a well-known approach in the discrete topology optimization.Ground structure is composed of uniform spaced nodes connected with each other by boundary conditions and external loads or forces. The ground truss structure is thought to encompass the potential optimal structure. Introductions to ground structure approach can be found in (Topping, 1993; Achtziger, 1997; Bendsøe and Sigmund, 2002).

Node positions in a ground structure are to be optimized as well as topology and truss bar cross sectional size optimization. This further node positions optimization is called geometric approach (Topping, 1993). In geometric approach, the node coordinates are also considered to be variables as well as bar cross sectional size. Research on both topology and geometry of ground structures can be found in (Ben-Tal, et al., 1993;( 2nO, where n is the number of the Achtziger, 2007). The complexity of ground structure approach is nodes. When n is large, the ground structure is very dense; and the numerical computation of LP or nonlinear LP techniques is unstable; and some unreasonable structures are obtained (Burns, 2002).)Rozvany, et al. (1995b) noticed that many literatures focused on truss structures rather than beam structures. They also pointed out that the beam structures are more practical than truss structures.

(1995b) noticed that many literatures focused on truss structures rather than beam structures. They also pointed out that the beam structures are more practical than truss structures. For truss structures, pin-jointed by only considering the axial forces and buckling is unavoidable. Beam structures are rigid-jointed and considering the shear forces which are matched with the engineering purpose. Theories on beam structure optimization can be found in (Rozvany, 1994; Rozvany and Prager, 1976). However, these theories are targeted for the continuum or partial continuum beam structure optimization. They cannot be simply applied in numerical computation. Some people combined the ground structure approach with the continuum based material optimization method to design the beam structures with joints (Fredricson, et al., 2003; Fredricson, et al., 2003; Fredricson, 2005; Kim, et al., 2008). These beam structures with joints are optimized with their sizes and can be manufactured directly.

These beam structures with joints are optimized with their sizes and can be manufactured directly. They are called compliant assemblies. Basically their physical model is a 3D solid one. Either sensitivity analysis or SIMP (simple isotropic material with penalization (Bendsøe and Sigmund, 1999) are utilized to optimize the size of each beam. Their optimization methods are continuum based.Motivated by simplifying the complexity of ground structure, we proposed a principal stress line method for the beam structure design. For the purpose, we also extended uniform strain energy density from truss structure to beam structure topology optimization. The remainder of the paper is organized as follows. Section 3 introduces the topology optimization problem and related mathematical formulations. The principal stress lines and their properties are also discussed in the section. Section 4 presents an overview of the principal stress line method.

The principal stress lines and their properties are also discussed in the section. Section 4 presents an overview of the principal stress line method. Details of the method are then presented in the following sections. Section 5 discusses the numerical method to compute the principal stress lines of a given design domain. Section 6 discusses the initial structure generation for a given design domain. Section 7 presents our size optimization method based on the uniform strain energy density principle. Section 8 discusses the topology growth of a candidate beam structure based on the principal stress lines. A test example is presented in Section 9. Finally Section 10 concludes the paper.3. PROBLEM FORMULATIONS AND OUR APPROACH3.1.

A test example is presented in Section 9. Finally Section 10 concludes the paper.3. PROBLEM FORMULATIONS AND OUR APPROACH3.1. Problem Formulation of Beam Optimization The optimization problem considered in the paper is to minimize the compliance on the total structural volume of a beam structure under static loads and constraints with the cross-sectional areas, the nodal coordinates, and the beam connections as design variables. Hence all the three types of optimization problems are considered including (1) size optimization by changing cross-sectional areas; (2) shape optimization by changing the nodal coordinates; and (3) topology optimization by changing the beam connections.

Even though most discrete optimization literatures consider truss structures, we focus on beam structures since they are more appropriate for SFF.Based on the linear elasticity theory (Eschenauer and Olhoff, 2001), the beam structure in our1 2 where u, K an F are the global displacement vector, the stiffness matrix and the load vector. Hence, the problem of minimizing the work done by external forces at equilibrium can be formulated as:problem follows the basic equations:TT(3.1)=∏WU −=uuFKu −,min χ B ∈Wu subject to ( )Ku =,F(3.2)Where B is a specific volume of body given byB={∫ χ χ Ωdx V =fixed}, and χ denotes the designvariable. By applying the equationU=1 2Wat equilibrium, the given optimum problem can also bealternatively written as:min Bχ∈uU )(subject toKu =,F(3.3)3.2.

By applying the equationU=1 2Wat equilibrium, the given optimum problem can also bealternatively written as:min Bχ∈uU )(subject toKu =,F(3.3)3.2. Michell Problem and Principal Stress LinesMichell (1904) first studies the fundamental properties of optimal grid-like continua and contribute exact analytical solutions for several well-known truss structures. One of them is Michell-cantilever, which is shown in Figure 2. Modern layout theory was founded by Prager and Rozvany (Prager and Rozvany 1977). Rozvany, et al. (1995a) also state that Michell trusses have the maximum stiffness for a given volume. According to Strang and Kohn (1983), the Michell problem is defined as follows:0=σdivin Ω ,n =:σfon Γ (3.4)Figure 2. Least-weight truss for a cntilever with a point load (Rozvany 1998).Hence, one can imagine that the bars of a truss-like continuum are placed in the directions of principal stresses.

Hence we havetan=2 θγ xy − εε yx(3.7)Therefore the principal stress lines are defined as the two orthogonal families of curves whosedirections at every point coincide with those of the minimum shear stress in a state of plane strain.3.3.

Principal Stress Line Analysis for Structural OptimizationFrom the conditions of elastic equilibrium, given the design domain, external forces and boundary conditions, in a finite element system of isotropic material, the following properties hold:Property 3.1 The displacement vector u is proportional to the external force vector F .Property 3.2 The displacement vector u is inversely proportional to the stiffness matrix K or the elastic modulus E.Property 3.3 The direction of the principal stress is not related to the scaling of the external forces, nor the material type for an isotropic material within the range of elastic deformation.Property 3.4 The principal stress field is mainly related to the topologic variables of the given structural design such as the position of external forces and the types of constraints.As an example, we present the principal stress field of a simple cantilever beam as shown in Figure 3.

For a single load with various sizes and different material properties, the principal stress lines arevirtually the same. This suggests that their optimal topology and shape design should be identical. Hence both load sizes and material properties will not affect topology design results. We also present the principal stress field for the same cantilever beam with different load or constraint positions. As shown in Figure 4.(a), the external force F is moved from the middle to the bottom of the right edge. As shown in Figure 4.(b), two point constraints are used instead of a line constraints in Figure 3. Since the principal stress fields have been changed, their optimal topology and shape design should also be different.(a) Steel, E=2.0e11[Pa], F=10[N](b) Aluminium, E=7.0e10[Pa], F=100[N]Figure 3. The principal stress fields varied loads and material properties.(a) Different load position(b) Different types of constraintsFigure 4.

The principal stress fields varied loads and material properties.(a) Different load position(b) Different types of constraintsFigure 4. Principal stress lines distribution in the continuum design domain.4. THE PRINCIPAL STRESS LINE METHODMotivated by Michell structures and the properties of principal stress lines, we developed a principal stress line method for designing a beam structure with the minimum compliance and maximum stiffness under any given loads and constraints. The optimization of size, topology, and shape of beam structure are all considered in our method. The design process of our method is shown in Figure 5 by using a single loaded Michell structure as illustration.

The design process of our method is shown in Figure 5 by using a single loaded Michell structure as illustration. There are five major steps including: (1) design domain specification, (2) principal stress lines computation, (3) initial structure generation, (4) size optimization, and (5) topology growth.For a given design domain specified by a user, the finite element method (FEM) is used to analyze the design domain based on the specified loads and constraints, the material information, and the maximum space that the resulting structure can occupy. Accordingly the principal stress lines can be computed and visualized. An initial structure is generated to connect all the constraints and loads. Based on the initial structure, size optimization is then carried out, which modifies the cross section area of eachbeam for improved performance. More principal stress lines can be added in topology growth, which will accordingly add more nodes and beams to the beam structure.

More principal stress lines can be added in topology growth, which will accordingly add more nodes and beams to the beam structure. Hence a new beam structure with different topology and shape will be generated for size optimization. The process is repeated until no performance improvement is found or a limit on number of beams has been reached. Sections 5~8 will explain in detail each of the five steps in the above process.Design domainInitial structure generationSection 6Section 7Size optimizationSection 5FFEM run with constant cross section areasFEM run with varied cross section areasCompute and visualize principal stress linesIterations end?NoVariation of cross section areas and scale to theYesgiven volumeFSection 8Topology growthNoResult OK?YesProposed Structure DesignFFigure 5. An overview of the principal stress line method for beam structure design.5.

An overview of the principal stress line method for beam structure design.5. PRINCIPAL STRESS LINE COMPUTATIONHegemier and Prager (Hegemier and Prager, 1969) utilizes the mathematical equations to compute the Michell truss bars based on the analytical computation of the directions of principal strains for a single loaded truss structure. The similar complementary slip lines are computed analytically in (Hill, 1950). In this paper, we develop a general numerical computation method to compute the directions of the principal stress lines. In addition to simple cases with a single load such as Michell truss bars, our numerical computation method can handle general cases with more complex boundary conditions and various external loads.5.1. Computing Principal Stresses at Any PointThe given design domain with loads and constraints is first analyzed based on the finite element method.

Computing Principal Stresses at Any PointThe given design domain with loads and constraints is first analyzed based on the finite element method. As shown in (3.6) and (3.7), the sizes and directions of the principal stresses are well defined for any given point in the stress field of a continuum 2D solid. However, in the 2D plane static analysis results generated by a FEM system (we used COMSOL3.2: www.comsol.com), the finite elements are a set of yσ and shear stress xyτ correspond to triangles and all the post-processing data of normal stresses triangle vertices. In order to query the post-processing data in any position in the design space, we compute the stresses based on linear interpolation. That is, suppose the stresses at position P is needed and assume P is located in a triangle element xσ , . The corresponding post-processing data (i.e. 3D separately.

That is, suppose the stresses at position P is needed and assume P is located in a triangle element xσ , . The corresponding post-processing data (i.e. 3D separately. The post-processing data D at the yσ or wD position P can be computed as: 3xσ ,21 PPPΔ 3xyτ ) for1P ,2P and3P are D =1D , 2D and uD vD + + 1(5.1)2whereu=→ PP→ PP 122××→ PP→ PP 133,v=→ PP→ PP 132××→ PP→ PP 113,w=→ PP→ PP 112××→ PP2→ PP 13. Hence the principal stresses1σ ,2σ and thecorresponding direction parameterized by θ at P can be computed.iP(cid:71) iv1+iPiP1+iP1−iP(cid:71) 1−ivδ1−iPFigure 6. Principal stress line search based on stress field points.5.2. Generating 2D Principal Stress Lines in a Continuum DomainBased on the principal stresses computed for a given point in the design domain, we will have four principal directions (i.e., in θ, 90 θ+(cid:68) ).

The closest direction selection(cid:71) 0, vP 0are the starting position and the principal direction; δis the search step;Whereis one of the(cid:71) 1−iviP closest to the search directionfour principal directions at position(cid:71) 1−ivis finished by comparing the angle with the search direction.Sufficient number of starting points is needed for generating principal stress lines that are dense enough to cover the entire design domain. We used two strategies in generating principal stress lines that are uniformly distributed in the design space.(1) For a simple design space, we add a guide curve and sample the curve uniformly. The samplingpoints are used as the starting points to generate the principal stress lines.(2) For a complex design space, we used the design space boundary as the guide curve and sample the curve uniformly. The sampling points are used as the starting points to generate the principal stress lines.6.

The sampling points are used as the starting points to generate the principal stress lines.6. DOMAIN SPECIFICATION AND INITIAL STRUCTURE GENERATIONAs shown in Figure 5, our method starts from the user specifying the loads and constraints, and also the maximum domain space that the resulting structure can occupy. Based on the computed principal stress lines, the user can identify the principal stress lines that connect the constraints and loads. We called such lines skeleton principal stress lines. An initial structure can then be generated from the skeleton principal stress lines. The generation of the initial structure can be performed automatically or by the user. For example, Figure 7.left shows the design domain for the Michell cantilever. The physics properties we used in the test are: H = 1.0 m; F = 1000 N; Young’s modulus E = 2.0e11 Pa; Poisson’s ratio ν= 0.33; /kg m ; Shape of a beam’s cross section = square; Given volume of the material = 0.015 Density = 7850 3m.

Figure 7.right shows the computed principal stress lines. The skeleton principal stress lines are also shown that connect the load and constraints. Accordingly an initial structure can be generated.3Figure 7. The design domain and initial structure of a Michell cantilever beam case.Notice after a candidate principal stress line is selected, we use a straight line connecting its two nodes in the beam structure instead of approximating the exact stress line by a set of polylines. This is because, for two given nodes in a beam structure, a straight line connecting them will lead to a smaller compliance of the structure than any other curves to connect them.7. SIZE OPTIMIZATION OF BEAM STRUCTURESIn our beam structure design, we follow energy principles which present the basis for both topology and size optimization of discrete and continuum structures.7.1.

Theoretical Basis – Axiom of Uniform Strain Energy DensityBased on the assumption of the positive definiteness of the global stiffness matrix, Achtziger (Achtziger, 1997) proved that the strain energy density is uniform among the truss structure. However, no references were found on beam structures. Here we will extend some conclusions of topology optimization of truss structures to the topology optimization of beam structures.As shown in (Achtziger, 1997), for the maximum stiffness topology optimization of loaded trussstructure, the optimization problem (3.2) can be reformulated as:V 2 Where f is a vector consisting of the external loads; u is the displacement vector consisting of all displacement components for all bars i and iK is the stiffness matrix of the i th bar. The proof can be extended to the minimization compliance or maximization stiffness topology optimization of discrete beam structures.

We also use the COMSOL script language to compute the strain energy of a beam structure. After a beam structure is analyzed, the strain energy is computed and saved in post-processing data. The strain energy density is expressed as strainenergy per unit length of a beam in COMSOL. As shown in Section 7.1, the objective in size optimization is to achieve the maximal stiffness of the beam structure. Based on the method of uniform strain energy density, to maximize stiffness also means to minimize the strain energy of the beam structures. In addition, the beam structure achieves the minimal potential energy when the strain energy density is uniform in the entire beam structure.We define the strain energy density as strain energy per unit volume of a beam, denoted by κwith / mJ unit of.

Hence we can optimize the cross section size of each beam for a given beam structure based on a process that is similar to the Fully Stressed Design Stress-Ratio Method in truss structures (Topping, 1993). However, different criteria are used. That is, instead of stress ratio, we use the strain energy density ratio to modify the cross section area of beam members after each stress/strain analysis of31 2⎛ κ ⎜ i ⎜ }{max ⎝ 1 ≤≤ After cross section areas of all beams are resized by the strain energy density ratio method, all cross VX (note zero cross section beams still have the section areas of beams are scaled to fit the given volume threshold cross section area ratio). With the resized beam structure, a new FEM stress/strain calculation is carried out, in which the cross section area in the bigger strain energy density beams gets bigger and bigger( in a relative way), and the cross section area in the smaller strain energy density beams gets smaller and smaller.

The process repeats until the strain energy distribution in the whole beam structure is close to uniform. An example of size optimization based on the above process is shown in Figure 8.⎞ ⎟ ⎟ ⎠beam structures. In each iteration, we use:(7.3)A=A⋅inewioldκ imi(a) A beam structure with uniform cross section; (b) resulted beam structure after size optimizationFigure 8: An test example of size optimization based on uniform strain energy density.8. TOPOLOGY GROWTH BASED ON PRINCIPAL STRESS LINESThe theoretical Michell structure has infinite number of bars and each bar is infinitely small. Hence such exact analytical solution is not practical for a beam structure design. Instead we would like to balance the number of beams and the performance of beam structure. For the purpose, we present a topology growth method based on the principal stress lines.

Instead we would like to balance the number of beams and the performance of beam structure. For the purpose, we present a topology growth method based on the principal stress lines. From an initial structure, the main process of our method is:(1) Identify new principal stress lines that can reduce the approximation errors between the beamstructure and the principal stress lines the most;(2) Use the identified principal stress lines to compute a set of intersection points as the positions ofinserted nodes (shape optimization);(3) Construct a beam structure by connecting the computed nodes following the connection of theprincipal stress lines (topology optimization).(4) Optimize the cross section size of each beam for the constructed beam structure (size optimization).

The strain energy of the whole beam structure can be computed, which can measure the stiffness of the beam structure.The above process repeats until desired performance is reached or the maximum number of beams is exceeded. This is illustrated in Figure 9 based on a Michell cantilever beam as shown in Figure 7. As shown in Figure 9, based on the initial structure in iteration 0, the maximum errors from the principal stress lines and the related beams are 1aP and 1bP . Accordingly two new principal stress lines are added which will lead to a new node 1cP . Therefore, a new beam structure can be constructed in iteration 1.After size optimization, the strain energy of the new beam structure is 3.61 mJ, which is significantly reduced from that of the beam structure in the previous iteration (4.15 mJ). Hence the stiffness of the beam structure has been increased for the same amount of material. The process can be repeated with more nodes and beams added.

Hence the stiffness of the beam structure has been increased for the same amount of material. The process can be repeated with more nodes and beams added. In the above topology growth process, it can be noticed that as more principal stress lines are used, the beam structure is closer to the Michell structure. The strain energy of the entire structure is getting smaller. However, the performance improvements are getting smaller after several iterations.Topology and shape of beam structuresBeam structures after size optimizationP1aIteration 0P1cP1bStrain Energy = 4.149 mJP2aP2dIteration 1P2cP2eP2bStrain Energy = 3.608 mJP3aP3dIteration 2P3cP3eP3bStrain Energy = 3.599 mJIteration 3Strain Energy = 3.530 mJFigure 9.

The topology growth of a Michell cantilever beam structure in 4 iterations.In our beam structure design process, we use the principal stress line as a bridge between the continuum design domain and the discrete beam structure design. This is different from other methods such as Homogenization, Ground Structure, and other expansion and growth methods for truss structure optimization (Kirsch, 1997; Martinez 2009). Compared to them, our method is dramatically faster. In addition, our method is easier to control. Hence the user can easily terminate or add more beams in the designed beam structure. However, a limitation of our method is that it can only be used for the minimum compliance or maximum stiffness structural optimization problem.9. TEST EXAMPLE: BRIDGE STRUCTURETo demonstrate the presented principal stress line method, another example of a bridge structure is presented.

TEST EXAMPLE: BRIDGE STRUCTURETo demonstrate the presented principal stress line method, another example of a bridge structure is presented. Similar to the Michell cantilever beam, the bridge structure is also a classical topology optimization test case. The design domain of a bridge structure is shown in Figure 10.left. The domain is fixed at the two end points of the bottom and a concentrated force is added in the middle of the bottom edge. The principal stress lines of the design domain based on the method discussed in section 5 are shown in Figure 10.right.Figure 10. a bridge like design domain and the principal stress lines.Based on the computed principal stress lines, the skeleton principal stress lines to connect the constraints and loads are shown in Figure 10.right. Notice the principal stress lines at two constraints have 45 degrees tangent with the horizontal axis. Accordingly, an initial structure can be generated which is also shown in the figure.

Notice the principal stress lines at two constraints have 45 degrees tangent with the horizontal axis. Accordingly, an initial structure can be generated which is also shown in the figure. The topology growth process of the structure is shown in Figure 11. After three iterations, the strain energy of the beam structure decreases from 1.33 mJ to 1.11 mJ. Hence the stiffness of the beam structure has been increased for the same amount of material used in the structure.Figure 11. The topology growth of a bridge structure in 3 iterations.10. CONCLUSIONSFor the minimum compliance design of beam structures, we presented a new structural optimization method based on principal stress lines. The mathematical basis of our discrete beam structureoptimization method has been presented. We also presented a numerical method for computing the principal stress field of any given design domain. Accordingly a topology growth method for refining beam structures has been presented.

We also presented a numerical method for computing the principal stress field of any given design domain. Accordingly a topology growth method for refining beam structures has been presented. In addition, based on the uniform strain energy density proposition in discrete truss structure, we extended the axiom of strain energy density to the discrete beam structure. The axiom of strain energy density is used in our size optimization of beam structures. Three examples were given to illustrate the design process of the principal stress line method. The results have demonstrated the effectiveness of our method. Compared to the ground structure method, the principal stress line method is much faster and easier to control. It is also more predicable and can overcome the numerical computation instability and results of unreasonable structures when the ground structure is highly dense.ACKNOWLEDGEMENTThis work is partially supported by the James H.

48(2), pp. 41-119, Feb. 1995a.30. Rozvany, G.I.N., "What is meaningful in topology design? An engineer’s viewpoint," Advances in structural optimization, pp. 149-185, Kluwer Academic Publishers, Dordrecht, 1995b.31. Rozvany, G.I.N., "Exact analytical solutions for some popular benchmark problems in topology optimization," Structural and Multidisciplinary Optimization, Vol. 15, pp. 42–48, Feb. 1998.32. Strang, G. and R.V. Kohn, "Hencky-Prandtl nets and constrained Michell trusses," Computer methods in Applied Mechanics and Engineering, Vol. 36, pp. 207-222, 1983.33. Svanberg, K., "Optimal truss sizing based on explicit Taylor series expansions," Structural and Multidisciplinary Optimization, Vol. 2, pp. 153-162, 1990.34. Svanberg, K., "Global convergence of the stress ratio method for truss sizing," Structural and Multidisciplinary Optimization, Vol. 8, pp. 60-68, 1994.35. Topping, B.H.V. "Topology Design of Discrete Structures," Topology Design of Structures, pp.

NCHMARKING EVALUATION OF AN OPEN SOURCE FUSED DEPOSITION MODELING ADDITIVE MANUFACTURING SYSTEMW. M. Johnson, M. Rowell, B. Deason, and M. EubanksEngineering Studies, Armstrong Atlantic State University, Savannah, GA 31419-1997AbstractThe availability of more affordable open source Additive Manufacturing (AM) systems has lead to the increased awareness and use of AM technologies. However, further expansion will necessitate improved reliability and an increased understanding in the limitations of these systems. This paper will review previous benchmarking models, and present the development of a new benchmarking model and its application in the evaluation of an open source AM system based on fused deposition modeling (FDM). The proposed benchmarking model includes various geometric features to evaluate the AM system in terms of dimensional accuracy, thermal warpage, staircase effect, and geometric and dimensional tolerances.1.

IntroductionAdditive manufacturing (AM) is a family of processes in which a numerical representation ofan object is used to fabricate the object one layer at a time. The AM processes are distinguished by the build materials and means by which the layers are joined together. Kulkarni et al. [1] provides three general classifications based on how the layers are joined: chemical bonding, sintering, and gluing. Fused Deposition Modeling (FDM) is an example of an AM process based on the intra and interlayer chemical bonding resulting from the extrusion of molten acrylonitrile butadiene styrene (ABS) thermoplastic from a heated nozzle. Stratasys Inc. first commercialized FDM in 1992 and continues to provide various FDM based systems.

Stratasys Inc. first commercialized FDM in 1992 and continues to provide various FDM based systems. However, their systems are proprietary (closed) with limited opportunities for use as a research platform in the enhancement of the FDM process.Since the middle of the last decade, open source and relatively inexpensive AM systems based on FDM have become available. These systems include the Fab@Home 3D printer [2], the RepRap project [3], and most recently, the CupCake CNC and Thing-O-Matic 3D printers by MakerBot Inc [4]. The Fab@Home 3D printer was originally designed with a non-heated syringe based extrusion system with thermosetting polymer for the build material, but can now be configured to extrude melted ABS plastic like the RepRap and MakerBot systems. The objective of the RepRap project was to develop an open source 3D printer that could be used to fabricate its own components (“self-replicate”).

The objective of the RepRap project was to develop an open source 3D printer that could be used to fabricate its own components (“self-replicate”). The MakerBot AM systems are derived from the RepRap project, but are more so focused on the fabrication of generic objects and have received quite a bit of general media attention [5-8], but has received limited evaluation and application in the literature. Pei et al. [9] recently studied the three previously mentioned open source AM systems and evaluated the capacity of a RepRap based system (Rapman) to fabricate geometricallycomplex parts. Kesner and Howe [9] mentioned the potential for both the Fab@Home and MakerBot systems for fabricating application specific components for force sensors.

Kesner and Howe [9] mentioned the potential for both the Fab@Home and MakerBot systems for fabricating application specific components for force sensors. Finally, Ludec-Mills and Eisenburg [11] noted both the Fab@Home and RepRap project, but they specifically utilized the CupCake CNC to fabricate parts based on 3D computer models created using their new spatial input device to introduce children to 3D modeling, design and fabrication. The CupCake CNC system, along with its other open source AM system counterparts are assembled by the end user.

The CupCake CNC system, along with its other open source AM system counterparts are assembled by the end user. Thus, there is some inherent variability and uncertainty in the limitations and overall performance of each CupCake CNC device.The objectives of this project are to review previous benchmarking models in the literatureused to evaluate AM systems, and to develop a new benchmarking model to evaluate the MakerBot CupCake CNC AM system in terms of dimensional accuracy, thermal warpage, staircase effect, and geometric and dimensional tolerances (GD&T). The outcome will provide the AM community with, to the authors’ knowledge, the first benchmarking evaluation of an open source FDM AM system. It will also help to expand the awareness and use of such systems as a research and development platform in the continued study of FDM AM process improvements and applications.

It will also help to expand the awareness and use of such systems as a research and development platform in the continued study of FDM AM process improvements and applications. It should be noted that this study is based solely on the configuration of our device and presents the corresponding findings of such.In what follows, details of the CupCake CNC will be provided, followed with a discussion ofbenchmarking models used in previous AM studies and of a new benchmarking model developed specifically for open source AM systems such as the MakerBot CupCake CNC. The paper will close with evaluation results of the fabricated benchmarking model and conclusions and about the MakerBot CupCake CNC system performance.2. Background2.1.

The paper will close with evaluation results of the fabricated benchmarking model and conclusions and about the MakerBot CupCake CNC system performance.2. Background2.1. MakerBot CupCake CNCThe CupCake CNC MakerBot (see Figure 1) consists of a horizontally translating (x-y direction) heated build platform (HBP) with a useable build area of approximately 80mm x 100mm and a vertically translating (z direction) extruder nozzle with a maximum build height of 130 mm. Positioning for each axis is accomplished using a belt driven pulley system with stepper motors. The extrusion system is also driven using a stepper motor to feed 3mm diameter ABS (plastic) filament into a heated 0.4mm diameter nozzle. The configured CupCake CNC used in this study can fabricate parts with layer thicknesses at 0.25mm, 0.31mm, or 0.36mm and does not include the capacity to fabricate support structures. CupCake CNC has an indicated accuracy and minimum feature size of 0.08mm [12].

CupCake CNC has an indicated accuracy and minimum feature size of 0.08mm [12]. The device is controlled via USB connection to a PC running the ReplicatorG open source software which includes Skeinforge. Skeinforge is a fully configurable Python based tool chain used to slice and output tool path instructions in G-Code format readable by MakerBot and RepRap. Examples of configurable process parameters within Skeinforge are detailed in Table 1. The road width (width of extruded filament) can bedetermined by multiplying the layer thickness and road width over thickness ratio which results in a road width of 0.6mm using the values from Table 1.Figure 1. Assembled CupCake CNC MakerBot.Table 1. Selected Skeinforge profile (fabrication) parameters.Skeinforge includes the option to fabricate objects using a raft structure, which serves as a platform upon which the object layers are placed.

Selected Skeinforge profile (fabrication) parameters.Skeinforge includes the option to fabricate objects using a raft structure, which serves as a platform upon which the object layers are placed. The raft helps to smooth any surface irregularities in the build platform and aids in the removal of the part from platform once fabrication is completed. As shown in Figure 2, the raft consists of evenly spaced parallel roads.The road spacing is intended to facilitate the removal of the raft from the fabricated part during post-processing.Figure 2. Example of the raft structure and a contour infill (infill solidity ratio of 0.40) of an object.2.2. Benchmarking Models in AM EvaluationIt is observed in the literature that benchmarking models are often used to characterize and evaluate the performance of both subtractive and additive manufacturing systems. Mahesh [13] identifies three types of benchmarking models as it pertains to AM: Geometric, Mechanical, and Process.

Mahesh [13] identifies three types of benchmarking models as it pertains to AM: Geometric, Mechanical, and Process. Geometric benchmarking models provide a means for system evaluation in terms of geometric accuracy, capacity to fabricate specific geometry, repeatability, and surface roughness. Mechanical benchmarking models can be used to determine and evaluate various material and mechanical properties of parts produced by AM systems, such as tensile strength and impact toughness [14-16]. Process benchmarking models can be used as a means to evaluate and/or optimize the AM fabrication process itself. Evaluation metrics for process benchmarking models might include those used in geometric [17-19] or mechanical benchmarking models, and also fabrication time and cost [20].

Evaluation metrics for process benchmarking models might include those used in geometric [17-19] or mechanical benchmarking models, and also fabrication time and cost [20]. Considering the potential for overlap across these three model types, a hybrid benchmarking model could be considered for a fourth type of classification in which the model incorporates various aspects of the geometric, mechanical, and/or process benchmarking models and associated evaluation metrics [21], [22].Benchmarking models have been used extensively in the performance evaluation and optimization of AM systems. As early as 1991, Kruth [23] evaluated the performance and possible applications of four AM processes (Stereolithography (SLA), Selective Laser Sintering (SLS), Laminated Object Manufacturing (LOM) and FDM using a geometric benchmarking model consisting of overhangs, an inclined cylinder, and various peg structures.

Juster and Childs [24] devised a more complex geometric benchmarking model to assess the performance of SLA, SLS, LOM, and FDM AM systems in terms of dimensional accuracy, geometrictolerances, repeatability, and fabrication limitations. The model included features such as square bosses, overhangs, freeform geometry, and vertical and horizontal holes.In 2000, Xu et al. [22] presented a benchmarking model to evaluate the horizontal (x-yplane) dimensional accuracy of the same set of AM systems examined by Juster and Childs. Given the primary intent of Xu’s study, the model consisted of various sized thin walls, notches, cylindrical bosses and an overhang structure. Mahesh et al. [25] used a benchmarking model previously developed by his co-author to evaluate four AM systems and processes.

Mahesh et al. [25] used a benchmarking model previously developed by his co-author to evaluate four AM systems and processes. The model incorporated more features than any other previous study and examined the AM systems’ performance based on dimensional accuracy, geometric and dimensioning tolerances and thermal warpage. Scaravetti et al. [17] and Campanelli et al. [18] both sought to improve the SLA process using process benchmarking models. Scaravetti et al. achieved limited success in their attempts to decouple observed dimensional and geometric (form and orientation) tolerances from the fabrication material versus the SLA device itself. Campanelli et al.

Campanelli et al. determined the optimal SLA process parameters (layer thickness, hatch overcure, and border overcure) to minimize the measured dimensional, positional and form deviations of the part relative to the CAD model.Overall, most benchmarking models and studies have been limited by small sample sizes,and mostly of the geometric benchmarking model classification to evaluate multiple AM processes in terms of geometric accuracy, surface roughness, repeatability, geometrical tolerances. The models have typically consisted of a rectangular base with various features to achieve the desired performance and evaluation metrics. These features included square and cylindrical bosses, vertical and horizontal through holes, inclines, thin walls, notches, overhangs, freeform geometry, fillets, and chamfers. The models were evaluated using coordinate measurement machines (CMM).

The models were evaluated using coordinate measurement machines (CMM). Based upon these previous works, a new geometric benchmarking model for the evaluation of the CupCake CNC AM system is presented. To the authors’ knowledge, this is the first benchmarking model to be specifically designed and presented in the literature for the evaluation of an open source AM system.3. Benchmarking Model DevelopmentBased on the benchmarking models discussed in the previous section, a new benchmarking model was designed to assess dimensional accuracy, feature size and geometry limitations (e.g., unsupported overhangs, and double curved geometry), geometric and dimensional tolerances, and repeatability all within the relatively smaller build envelope of the CupCake CNC. Table 2 and Figure 3 provide a brief description and visual representation (CAD model) of the specific features of the new benchmarking model.Table 2.

1, 1.5, & 2mm thicknesses. 3x10mm, 3mm deep cavitycylindricity, roundness, and repeatability angularity, sloping smoothness and linear accuracy profile, slope changes, and symmetrylinear accuracy, parallelism, and wall thicknessCircular holes C21, C6, C5Circular holes OverhangsC28 & C27Square notchesXN1.5-4 & YN1.5-45mm Dia, 3mm deep, 15mm Dia, 20mm Dia.(in base feat). roundness, relative position, and repeatability 5mm & 10mm Dia.3mm deep. (in tower feature) 40, 45, 50 degrees 1.5, 2, 2.5, 3, 3.5, & 4mm, spaced 5mm apart. (Qty 2)roundness, relative position, and repeatability angular accuracy and sloping limitlinear accuracy and consistencyFigure 3. CAD version of benchmarking model with referenced feature identification (IDs)4. Benchmarking Fabrication and Evaluation ResultsThe benchmarking model was fabricated with the parameters discussed in Table 1. Themodel took 2.5 hours to fabricate and used 36.4 cm3 of ABS filament.

Themodel took 2.5 hours to fabricate and used 36.4 cm3 of ABS filament. Figure 4 (a-c) presents the original CAD model along with angled front and back, and top views of the fabricated benchmarking model. The following observations were based on a visual inspection of the part: 1. Notches were fabricated, but did not have uniform gap spacing especially towards thebottom of the base feature.2. Holes in base appeared circular, but showed nonuniform interior surfaces. 3. Horizontal holes on the tower feature and tower base appeared uniform. 4. Warping of 45o and especially the 50o overhangs was present (Figure 4c). 5. Warping was also visible along the edges of the bottom of the base feature. 6. All vertical sides of the base feature tapered outward from the bottom surface. 7. All thin walls were fabricated, but the 1.5 and 2mm walls had core voids. Core voids were also present at two other locations on the top surface of the base feature. (See circled areas in Figure 4d)8.

Core voids were also present at two other locations on the top surface of the base feature. (See circled areas in Figure 4d)8. The staircase effect was inversely related to the inclines angles. The staircase effect was present to a lesser degree in the hemisphere.Figure 4 (a-d). CAD (a) and fabricated (b, c and d) versions of benchmarking model4.1. 3D Laser Scanning Results and DiscussionThe benchmarking model was scanned by a service provider using a 3D laser scanning system. A laser scanning approach is uniquely suited for the inspection of parts fabricated using AM considering the possibility of layer-wise deviations. However, 3D laser scanning systems are relatively less accurate than coordinate measurement machines (CMM) traditionally used for inspection. The 3D laser scanning system used had an accuracy of 0.0089mm and was suitable for our inspection given the stated 0.08mm minimum feature size limitation of the CupCake CNC.

The 3D laser scanning system used had an accuracy of 0.0089mm and was suitable for our inspection given the stated 0.08mm minimum feature size limitation of the CupCake CNC. Furthermore, our benchmarking model contains features no smaller than 1.0mm.Based on the scan of the fabricated model, a total of 2,999,732 data points were used to determine the deviation from the nominal CAD model dimensions after aligning the scan data with CAD model using a best fit approach. Deviations are defined based on the shortest distance between a scanned data point and any point on the CAD model. Figure 5 (a-d) provide the subsequent error (deviation) maps for various orientations of the model. The areas of the model in green are within +/- 0.1 mm of the CAD model, red areas are outside the CAD model (oversized) by 1mm or greater, and areas in blue are inside the CAD model (undersized) by - 1mm or less.

The standard deviation for the data is 0.3101mm, with 98.46% of the points within +/- 2 standard deviations, 79.2% of the points within +/- 0.2mm, and 97.7% of the points within +/- 0.5mm of the nominal CAD dimensions.Most of the observations from the visual inspection were readily confirmed in the error maps of Figure 5(a-d). For example, Figure 5c shows warping along the edges of the bottom surface. The rest of the bottom surface has mostly a positive deviation likely due to the incomplete removal of the raft structure. Observations 7 and 8 were not as apparent in the error maps due to the use of color averaging (color averaging does not influence the numerical results). The positive (red) deviation of the 25o incline (A12) was not obvious from the visual inspection. However, this deviation was a result of the inadvertent use of a benchmarking CAD model version with an A12 incline angle of 30o instead of 25o for the deviation study.

However, this deviation was a result of the inadvertent use of a benchmarking CAD model version with an A12 incline angle of 30o instead of 25o for the deviation study. All other incline features on the CAD model were verified and consistent with the value reported in Table 2 and Figure 3.In addition to the error map plots, the 3D scan data was also used to obtain the dimensions ofvarious features on the fabricated benchmarking model. A listing of additional dimensional values for the model features is provided in Appendix 1. Deviation values were computed using the CAD model dimensions as the reference. Unless otherwise noted, the average of three cross sections of the scan data were used to determine the mean dimensions of each feature. All dimensions and deviations are in millimeters. Table 3 shows the overall dimensions of the fabricated model.

All dimensions and deviations are in millimeters. Table 3 shows the overall dimensions of the fabricated model. The tower width had the largest deviation of the group which was expected given the warping of the 50o overhang evident in Figure 4b and c.As shown in Table 4, the cylindrical boss diameter deviations were all positive (oversized).The 6mm diameter cylinders had a range of 0.11mm. Based on Table 5A, the heights of the cylindrical bosses are defined along the z-axis and thus subject to an inherent deviation due to the fixed layer thickness (0.36mm). For example the 7 mm tall, 6mm diameter cylindrical boss would require 19.44 layers, which rounded down to 19 layers results in an expected negative deviation in the boss height. All boss height deviation values are consistent with being over or undersized based on the expected deviation due to layer thickness.Figure 5(a-d). 3D Comparisons (error maps) of benchmarking model based on nominal CADdimensions.Table 3.

3D Comparisons (error maps) of benchmarking model based on nominal CADdimensions.Table 3. Overall dimension and deviations of benchmarking model.Table 4. Cylindrical boss diameters and deviations.Feature Description 6mm Dia Boss,7mm tall C4 6mm Dia Boss,7mm tall C20 6mm Dia Boss,7mm tall C24 10mm Dia Boss, 5mm tall C23 20mm Dia Boss, 7mm tall C7IDCAD Dim.66610Mean Dim. Deviation6.260.266.150.156.210.2110.190.192020.170.17The hole diameter and deviations in the base feature (C5, C6, and C21) are presented in Table 5B. There was insufficient data to determine the C5 diameter, but the other two diameters were undersized (negative deviations). These results are at least in some aspects due to our Skeinforge settings which did not compensate for road (filament) width when traveling in circular tool paths (in the x-y plane). The “stretch” Skeinforge plugin can be used to account for this issue by essentially widening the hole diameter.

The “stretch” Skeinforge plugin can be used to account for this issue by essentially widening the hole diameter. Although the diameters for C27 and C28 were also undersized, they are instead (at least) affected by the tool path in the x-direction and the layer thickness.The notch feature (XN and YN) gap dimension deviations are plotted as a group in Figure 6. Four of the six notch gap (absolute) deviations where greater in the x-direction than in the y- direction, with the 2mm notch gap in the x-direction having the largest absolute deviation (0.35mm) of the group. All but one of the gaps in each direction was oversized (positive deviation). Mostly positive feature deviations were also noted for the larger single axis (x or y- direction) tool paths associated with the width and depth dimensions of the rectangular bosses, A1-3 and B1 and B2 (See Appendix A). The ranges for the width and depth dimension of the square bosses were respectively, 0.13mm and 0.19mm.

The ranges for the width and depth dimension of the square bosses were respectively, 0.13mm and 0.19mm. Notwithstanding the relatively small sample size, the consideration of the range values of the square and cylindrical bosses suggests some level of repeatability in the fabrication of these features.Figure 7 shows a cross sectional view of the scanned inclines for the 15o, 25o, and 35oinclines, with the red lines showing the CAD model incline. A screenshot of the 80o incline is not shown, but measured 77.4o. The staircase effect is reduced as the incline angle increases. This relationship was noted by Pandey et al. [26].Table 5. Heig T s. Holesght/depth vallues and assoociated deviations for A) Circular bo) osses and B)Table Featu Descre A. ure riptionD IDCAD Dim.Mean M Dim. Deviatio DonTable B. Feature DescriptiionIDCAD Dim.Mean Dim.

Th s with horiz and A2 feat erance zone tolerance zo top surfacening toleranc was compar 24) and throu e staircase e ontal axes (C tures were c due to the th ones were als of the base fce results are ed with the c ugh holes (C effect was a c C27 and C28 omparable, w hermal warp so comparab feature as the shown in T correspondin C6 and C5) h contributor t 8). The flatn with the top page of the 5 ble with the e he datum).Table 6. The ng CAD feat had reduced t to the cylind ness of the to surface of th 50o overhang exception ofe scanned da ture. The tolerance zo dricity toleran op surfaces o he tower hav g. The f the front toatanes nce of the vingwerFigure 7. Cross sectional views of the 150, 250, and 350 inclines. CAD inclines in red andscanned data in black.Table 6. Geometric and dimensioning tolerance results for selected features.Feature DescriptionID GD&T Feat.Meas. Tol. ZoneFeature DescriptionIDGD&T Feat.Meas. Tol.

Front Base- Datum Tower (front). Base (top)-Datum B2 (front). Base (top)- Datum B1 (right side). Tower Base (top)- Datum16152217Perpendicu larityPerpendicu larityPerpendicu larityPerpendicu larity0.9420.810.455. ConclusionA benchmarking model was developed to evaluate the performance of the MakerBot CupCake CNC AM system. The evaluation metrics included dimensional accuracy, staircase effect, and thermal warpage, and geometric and dimensional tolerances. The benchmarking model had a square base (70x70x10mm) and included features such as thin walls, rectangular and cylindrical bosses, through and blind holes, inclines, notches, and a hemisphere. The overall height was 56mm. The part was successfully fabricated with noticeable warping along the bottom edges of the base feature and for the 450 and 500 overhanging inclines (unsupported during fabrication).

The part was successfully fabricated with noticeable warping along the bottom edges of the base feature and for the 450 and 500 overhanging inclines (unsupported during fabrication). Based on 3D laser scanning of the part, most of the measured data points had positive deviations, and overall 97.7% of the points were within +/- 0.5mm of the nominal CAD dimensions. These results are based on a single sample produced with a configuration specific to our device. Notwithstanding these limitations, the results suggest that the MakerBot CupCake CNC is a viable alternative to more expensive propriety FDM AM systems.

Notwithstanding these limitations, the results suggest that the MakerBot CupCake CNC is a viable alternative to more expensive propriety FDM AM systems. Future work will evaluate the fabrication of the benchmarking model using a commercial FDM system and comparison of the results from the model fabricated with the CupCake CNC system.AcknowledgementsThe first author would like to thank the following students: Brian Redden, Matt Carroll and Michael Tam for their assistance in the preliminary configuration and assembly of the MakerBot system and literature searches. This work was supported with funds from the Armstrong/NASA Space Grant Program, the Armstrong/NSF-STEP (Science Technology Expansion Program, DUE-0856593) Grant, and a research and scholarship grant from the Armstrong Office of Sponsored Programs.Bibliography[1] P. Kulkarni, A. Marsan, and D. Dutta, “A review of process planning techniques in layeredmanufacturing,” Rapid Prototyping Journal, vol. 6, no. 1, pp.

NCHMARKING OF DIFFERENT POWDER-BED METAL FUSION PROCESSES FOR MACHINE SELECTION IN ADDITIVE MANUFACTURINGE. Yasa*, F. Demir*, G. Akbulut*, N.Cızıoğlu*, S. Pilatin*TUSAS Engine Industries, Inc. Eskişehir, TurkeyAbstractIn the last decade, additive manufacturing has gained significant interest for direct part production and started to change the way companies manufacture products; even in very demanding sectors like aerospace. The biggest challenge for a wider industrial acceptance still stands as the need for more reliable, repeatable and precise machines for additive manufacturing. This paper presents a comprehensive benchmarking study for the selection of an additive manufacturing machine for powder-bed metal fusion process, i.e. Selective Laser Melting or Direct Metal Laser Sintering or Laser Cusing.

Selective Laser Melting or Direct Metal Laser Sintering or Laser Cusing. Four different machine vendors for the same technology to be employed for aeroengine part manufacturing using Inconel 625 powder have been involved for comparing different machine specifications. Many aspects such as dimensional accuracy, surface quality, need of support structures, density, hardness and process limits (minimum wall thickness, overhang surfaces, inclinations and curvatures, etc.) are addressed in the paper.

The state-of-the-art in machines for powder-bed metal fusion process is presented aiming at understanding the current limitations of the technology available today.IntroductionLaser Beam Melting (LBM), also referred to as direct metal laser sintering/melting (DMLS/DMLM), Laser Cusing®, Selective Laser Melting® (SLM), is a powder-based, additive manufacturing process where a 3D part is produced, layer by layer, by using a high-energy lasers to selectively melt powdered materials which them fuse during solidification as defined in VDI 3404 (see Figure 1 for schematic LBM process). The process involves physical binding activated by the heating of the powder particles using the laser energy.

The process involves physical binding activated by the heating of the powder particles using the laser energy. Common post-processes to improve surface quality may include microblasting, laser re-melting or laser assisted material removal [1].Figure 1: Schematical Laser Beam Melting process [1]390Among many additive manufacturing (AM) technologies, the processes suitable and commercially available for functional metallic materials using electron or laser as the energy source are demonstrated in Table 1. There are other technologies targeting metals such as sheet lamination or binder jetting but these are less relevant in terms of direct functional part production, but rather prototyping [2]. Depending on many criteria such as material suitability, surface roughness, accuracy, component size, needed post-processing technologies, etc. one has to decide which AM technology is best for her/his application.

one has to decide which AM technology is best for her/his application. Choosing the needed AM technology is most of the times is not the final decision to be made (see Table 1). Among many machine vendors, the most suitable brand and model has to be decided, which is not always an easy process to accomplish. In case of laser beam melting, all machine vendors claim to manufacture the state-of-the-art machines and there is unfortunately no standard yet established for machine qualification or acceptance testing in the open literature. Companies in need of an LBM machine usually specify their own acceptance testing procedures or use the machine vendors’ own established qualification procuredures Not only companies to buy a machine, but also the research institutes carry out benchmarking studies to compare different processes and also different machines. Some examples of benchmark designs are given in Figure 2.

Some examples of benchmark designs are given in Figure 2. Various studies focus on determining the process capability in terms of different aspects. For example, Vandenbroucke and Kruth focused on identification of the process limits regarding the accuracy, surface roughness, density, mechanical properties and stair effect for two different materials, i.e. Ti64 and CoCr, for dental applications and the tests were carried out on one LBM machine, namely Concept Laser M3 machine. This study more aims at optimization of the process parameters for maximum part performance [8]. Kruth et al., in [9], compare different additive manufacturing technologies, i.e. Selective Laser Sintering -SLS and LBM, for evaluating these processes as rapid manufacturing technologies rather than rapid prototyping.

Selective Laser Sintering -SLS and LBM, for evaluating these processes as rapid manufacturing technologies rather than rapid prototyping. As a result of this study, they have concluded that the real breakthrough of SLS and LBM technologies in other industries than medical parts and tooling inserts, depend on further improvements in process accuracy and productivity [9]. Campanelli et al., in [10], also investigates the process capabilities and performance of the LBM process by testing density, accuracy and mechanical properties. In this study, it is concluded that the process is capable of producing almost full density parts with good mechanical properties with a good accuracy for nominal dimensions greater than 400 µm. For nominal dimensions lower than this value, the errors increase [10]. Although most of the studies target at exploring process capability/limits, some studies more focus on equipment comparison for a single process.

Although most of the studies target at exploring process capability/limits, some studies more focus on equipment comparison for a single process. Mantel, compares laser cusing, direct metal laser sintering and selective laser melting which are commercial names for the same process,i.e. LBM, respectively by Concept Laser, EOS and SLM Solutions [11]. Despite the fact that different materials were used for comparison for the same geometrical and material properties, different benchmark geometries were tested in this study concluding that Laser Cusing is the most promising LBM technology for different industrial sectors such as tooling, medical, aerospace and motor industry (see Figure 3 for comparison results for aerospace). However, taking the materials used in the study into account, that may be a misleading information. In this paper,the results of a benchmarking study carried out to select the most suitable LBM eqipment for aerospace applications are presented.

In this paper,the results of a benchmarking study carried out to select the most suitable LBM eqipment for aerospace applications are presented. The procedure was first to design a benchmarking part, which is actually a modified version of the one presented in [3] due to timing issues (see Figure 4 for the design used in this study). This benchmark part design was sent to four LBM machine vendors (Renishaw, Concept Laser GmbH, EOS GmbH Electro Optical Systems, and SLM Solutions GmbH) to build the parts in Inconel 625. It was mainly specified that 1) Support structures may be utilized for overhang surfaces but they should be removed before shipment; 2) Heat treatments391only for stress relief may be applied on the parts; 3) Material removal processes such as grinding, polishing, benching are not allowed on any parts.

After the parts were delivered by TEI, they were evaluated in terms of detail resolution, geometrical accuracy, surface quality, density and hardness comparing the results with each other.Table 1: Various AM technologies and commercially available machine vendors for direct functional metallic parts using electron or laser as the energy sourceabcdeFigure 2: Benchmark designs from different studies a) [3] b) [4] c) [5] d) [6] e) [7]392Figure 3: Benchmark test results regarding the aerospace and motor industry in [11]Figure 4: Benchmark design used in this studyBenchmark Geometrical EvaluationAs a starting poing, the part design were sent to all four machine vendors in .stp format and no specific direction for the build direction was given although in ASTM 52921 standard specifies z- axis as the build direction in its definition. Z-axis, for processes employing planar layerwise addition of material, shall run normal to the layers [12] (see Figure 5).

Z-axis, for processes employing planar layerwise addition of material, shall run normal to the layers [12] (see Figure 5). The received benchmark parts from four machine vendors are shown with related build directions in Figure 6. As shown in the figure, not all machine vendors used the same axis for the build direction. For manufacturing the inclined walls with holes on, two of the machine vendors (V1 and V4) chose to build the benchmark part along Y-axis. In this case, there was no need to support the inclined walls whereas the small pin and the pipe detail needed to be supported. Although no polishing was allowed on the parts, probably due to very bad surface quality on the curved surface, some polishing may have been applied for the case of V4 as depicted in Figure 7a whereas Figure 7b shows the difficulty of making the boss detail in the selected build direction. The machine vendors were also asked to send the parts without supports which was not the case for V3.

The machine vendors were also asked to send the parts without supports which was not the case for V3. Moreover, during cutting the part from the base plate, the thin section connection overhang inclined walls and the main body was lost for V3. The geometrical details of the benchmark parts are demonstrated in Figure 8 regarding sharpness of vertical walls, boss, top surface, overhang393inclined walls, pipe detail and others. The geometrical details of the benchmark were generally made successfully by all four vendors with a few exceptions despite the quality difference. The summary of the results in terms of geometrical features are as follows:The sharp edges with different angles were best manufactured by V1 and V3 although they used a different build direction (all build directions were shown on the pictures. If not, it is from the page upwards or downwards).

If not, it is from the page upwards or downwards). There is a lack of straightness for the sharp edges of V2 and V4 as shown in the first two rows of Figure 8.The boss detail is actually a tower of cylinders with decreasing radii as shown in the third row of Figure 8 (nominally from bottom to up 5, 2, 1 and 0.5 mm respectively). The smallest cylinder with a diameter of 0.5 mm was only successfully made by V3 with an accuracy of approximately 20 µm. All others failed to produce the smallest diameter properly. The complete boss of V4 is unacceptable due to bad surface on the overhang side due to teh build direction selection The cylindiricity of the boss made by V1 is also unacceptable due to the same reason. The surface quality and thus sharpness of the boss is best by V2 (see row 3&4 in Figure 8).The pipe detail is probably one of the most difficult features on the benchmark. The details of this feature on V3 is not clear due to supports.

The details of this feature on V3 is not clear due to supports. The selected build direction for V4 made it very difficult to produce this feature whereas V1 successfully made this feature along the same build direction with V4, probably due to good design of the supports. V2 has made it succesfully with the dross formation on the overhang surface with a good surface quality both inside and outside of the feature (see row 5 in Figure 8).Figure 5: Generic AM machine system and coordinates [12]b Figure 6: Benchmark parts from 4 different machine vendors with the utilized build directions1 a) top view b) side viewa1 Machine vendors’ names are on purpose not speficied394Build directionFigure 7: a) Traces of polishing b) overhang surfaces for the boss detail for V4The overhang inclined walls were successfully made by V1 and V4 due to the build direction. For V2 and V3, there was a need to support overhang walls with two maximum overhang angles.

For V2 and V3, there was a need to support overhang walls with two maximum overhang angles. The removal of all support structures was successful for V2 (see Figure 9) although some particles remained trapped in the root (see row 6 in Figure 8).The holes with different diameters (5, 2, 1 and 0.5 mm) were best made by V2 with an accuracy of approximately 100 µm whereas for V4 the accuracy for the hole diameters is around 200 µm. Due to the selected build direction, V1 and V4 benchmarks show a dross formation on one side of the holes for almost all of the diameters. Moreover, the holes with a diameter of 1 and 0.5 mm are partially blocked for V1. For V3, the hole with a diameter of 0.5 is totally blocked whereas the hole with a diameter of 1 mm is partially blocked due to partially melted powder particles sticking in the holes (see row 7 in Figure 8).The overhang surface on the side of the benchmark all seem to be made with supports for the rectangular hole (see row 8 in Figure 8).

However, the form of both features were not good for V1, V2 and V4 with accuracy values greater than 200 µm. For V3, the geometrical accuracy was less than 100 µm despite the support structure is still present for the rectangular geometry and makes it difficult to evaluate.The last observed feature is the thin walls with a nominal thickness of 250 µm. For V4, it was not possible to make the walls in the build direction properly. For V1, the same problem appeared but the horizontal thin wall was manufactured with a higher thickness (~580 µm) due to dross formation whereas the straightness of the vertical wall is not acceptable. The best thin walls were made with V2 and V3. However, the side surface quality of the thin walls was very bad for V3 while the accuracy is higher than the one of V2 (see Figure 10 for details and dimensions).

However, the side surface quality of the thin walls was very bad for V3 while the accuracy is higher than the one of V2 (see Figure 10 for details and dimensions). Surface Roughness EvaluationIn the scope of this study, only Ra and Rz, are taken into account for comparing four benchmark parts on two surfaces perpendicular to each other. As commonly known, average roughness, Ra, is the arithmetic mean of all deviations from the center line over the sampling path whereas Rz is the average distance between the five highest peaks and the five deepest valleys within the sampling length. The surface roughness in terms of average roughness (Ra) and surface roughness depth (Rz) are demonstrated in Figure 11 for four benchmark part with 95% confidence intervals.. The surface roughness results were taken on two surfaces of the benchmark part as shown in Figure 12.

The surface roughness results were taken on two surfaces of the benchmark part as shown in Figure 12. As shown, taking the build directions differently, the top and side3951 w o R2 w o R3 w o R4 w o R5 w o R6 w o R7 w o R8 w o R9 w o RV1V2V3Figure 8 Benchmark details396V4surface definitions differ for V1and V4. The measurements were taken with MarSurf M300 with a cutoff length of 2.5 mm, 3 repetitions and a sampling length of 12.5 mm. As Figure 11 depicts, the highest surface roughness (Ra and Rz) on the side surface is exhibited by V3 whereas the highest surface roughness (Ra and Rz) on the top surface is shown by V1. The trends in Ra and Rz are the same. V2 shows the lowest side surface roughness: Ra of 3.9 ± 1.4 µm and Rz of 24 ± 1 µm whereas V3 shows the lowest top surface roughness values: Ra of 7 ± 0.5 µm and Rz of 35 ± 3 µm. However, V3 shows 4 times higher roughness (Ra) on the side surfaces compared to V2.

However, V3 shows 4 times higher roughness (Ra) on the side surfaces compared to V2. Taking side and top surfaces into account, V2 and V4 gives the best and similar surface quality results.Figure 9: Remains from the support structures in the overhang inclined surfaces - V2a)b) Figure 10: Thin wall details of a) V2 and b) V3181614Ra120100Rz12801086TOPSIDE6040TOPSIDE422000V1V2V3V4V1V2V3V4Figure 11: Surface roughness measurements a) Average roughness Ra b) Surface roughness depth Rz397Figure 12: The surface roughness measurement directions for side and top surface notationsDensity EvaluationFor the evaluation of the density, two complementary methods were utilized: Archimedes’ method and image analysis. For the Archimedes’ method, the solid section, as shown in Figure 13 with red color, is cut from the benchmark part and divided into three equal parts.

For the Archimedes’ method, the solid section, as shown in Figure 13 with red color, is cut from the benchmark part and divided into three equal parts. The density of each part was measured by Archimedes’ method and 8,44 g/cm3 was taken as the reference density to calculate relative density for each benchmark part section [13]. Figure 14a shows the obtained results with 95% confidence intervals. V2 and V4 show densities above 99% whereas V3 has a relative density higher than 98%. Yet, V1 stays hardly above 97%. The results obtained from Archimedes’ method are more global than image processing and do not give a clue for reasoning low densities. Thus, the cross-sections of specimens from the same block were observed to investigate the reason of such low values. An in-house developed Matlab code was utilized to calculate the porosity in the cross-sections.

An in-house developed Matlab code was utilized to calculate the porosity in the cross-sections. 3 images for each benchmark part were used and the average porosity values are shown in the Figure 14b where porosity values stay below 0.3% for all benchmark parts. One of the used images for each benchmark part is depicted in Figure 15. The image for V1 shows a high density as opposed to the values obtained by Archimedes’ method. Some other sections from V1 were also measured by both methods and the results were the same. The reason may be a significant change of the chemical composition being different than others resulting in a lower density than 8,44 g/cm3. The material compositions have not yet been investigated to be sure about this probability up to now. It can be concluded that laser beam melting gives almost full density for V2 and V4 with no problems associated.

It can be concluded that laser beam melting gives almost full density for V2 and V4 with no problems associated. For V3, the lower density may a consequence of need of process parameter tuning or bad powder coating (see Figure 16 for traces).Figure 13: Section cut from the benchmark parts for density measurements398Relative Density by Archimedes' MethodV4V3V2V195%96%97%98%99%100%Relative Density by Image AnalysisV4V3V2V199,60%99,70%99,80%99,90%100,00%Figure 14: Relative density measurements by a) Archimedes’ method b) Cross-sectional image analysis methodHardness EvaluationSince hardness is a fast measure of many mechanical properties such as ductility, plasticity, strength and toughness, only hardness measurements were accomplished to get a quick and general idea of the mechanical properties obtained after LBM.

Thus, microhardness of four benchmark parts were measured using indentation Vickers hardness using 500 gr force with approximately 10 s. The microhardness measurments were preferred against macrohardness measurements so that the results represent the material hardness but not the porosity. The measurements were taken using the same specimens used for density measurements on surfaces parallel to the build direction for all benchmark parts. 25 measurements were taken on each benchmark specimen. The results are demonstrated in Figure 17 with 95% confidence intervals. The highest hardness value is obtained from V2 with minimum deviation among different measurements (333 ± 3 HV) whereas the lowest hardness was derived from V4 with 299 ±5 HV. The values above 300 HV are in accordance to the values given by EOS Inconel 625 datasheet and other studies conducted with additive manufacturing of Inconel 625 [14-16].

The values above 300 HV are in accordance to the values given by EOS Inconel 625 datasheet and other studies conducted with additive manufacturing of Inconel 625 [14-16]. Compared to the hardness obtained with solution strengthened Inconel 625 in wrought form and annealed399condition generally below 275 HV, laser beam melting leads to a higher hardness probably due to high cooling rates and fine microstructure [13].V1V2V3V4Figure 15: Cross section samples used for density analysisFigure 16: Waviness observed on the top surface of V3The microstructures of the benchmark parts are also shown in Figure 18 and Figure 19 for different planes. Although a more detailed study is needed to evaluate the microstructures obtained, it can be stated that V1 shows a very fine and different microstructure compared to others. This also suggests a chemical composition variation as is suggested by density measurements.

This also suggests a chemical composition variation as is suggested by density measurements. Figure 19 demonstrates a microstructure with fine dendrites and cells for both benchmark parts. Figure 19 also shows an apperant layering with curved melt scan bonding lines and the contrast provided to illusrate the the melt bonds is probably due to γ″ precipitates concentrating in these regions as also suggested by [17].400V1V3V H s s e n d r a HV H s s e n d r a H350 325 300 275 250051015 Measurement #202530340320300280V1V2V3V4Figure 17: Microhardness test resultsV2V4Figure 18: Microstructures in four benchmark parts (XY plane)401V1V2V3V4V4V3Figure 19: Apperant layering with curved melt scan bonding lines with the build directions shown for (XZ plane)ConclusionsA comprehensive benchmarking study for the selection of an additive manufacturing machine for laser beam melting process has been presented in this study.

Four different equipment providers for LBM to be employed for aeroengine part manufacturing using Inconel 625 powder have been involved for comparing different machine specifications. Many aspects such as dimensional accuracy, surface quality, geometrical resolution, need of support structures, density, hardness and process limits are addressed in the paper. To conclude, it can be stated that laser beam melting technology is getting more and more mature every day for many diverse industrial sectors. However, it can still not be said that it is a directly plug and play technology for functional part production for applications with high requirements as commonly mistaken in public due to the easiness of personal 3D printers. For each application, in terms of part and material properties, the process has to be developed and optimized for best performance.

For each application, in terms of part and material properties, the process has to be developed and optimized for best performance. Moreover, for the machine selectin, there are other factors than part properties, which should be taken into account such as equipment price, lead time, warranty time, consumables, optional functions or modules (data logging, powder coating or melt pool monitoring, processing in vacuum atmosphere, glass scale for linear drives, etc.) and so on. A decision matrix can be utilized to weigh each factor according to the specific needs.AcknowledgementsThe authors would like to thank their colleagues at TEI; Özgür Poyraz, Rifat Yılmaz, Ezgi Uğur Solakoğlu, Soner Ören and Aydın Yağmur for the technical support and Undersecretariat for Defense Industries of Turkey in the scope of YAKUT project.References1. VDI 3404 Additive Manufacturing Basics, definitions, processes, Draft version, Düsseldorf, May 2014.2.

nchmarking of Rapid Prototyping Systems - Beginning to Set StandardsDureen Jayaram, Amit Bagchi, C.C. Jara-Almonte, Clemson University Sean O'Reilly, Ford Motor CompanyAbstractMany rapid prototyping (RP) technologies are available today and more are being developed around the world. The absence of benchmarking standards in the RP industry has led manufacturers to use their own standards and make claims about superior performance. The need for testing standards is already felt; standardization will become imperative in the near future. The present work aims to lay the groundwork for the development of standards to measure various Issues such as appearance and finish are studied qualitatively; the test part performance factors. and some findings are presented. Issues such as repeatability, warpage, curl, creep, shrinkage and tensile strength are proposed to be studied quantitatively; test parts designed for studying these are described.

Issues such as repeatability, warpage, curl, creep, shrinkage and tensile strength are proposed to be studied quantitatively; test parts designed for studying these are described. Benchmarking standards will help users choose proper systems for their applications and help operators in monitoring machine performance, enabling better control over part building.1. IntroductionToday's various rapid prototyping (RP) systems have significant differences with respect to their operation and the materials they use. Experimental evidence shows that there are different factors which determine the geometries that build better on each process, the speed of building, the finish and strength of the parts and the accuracy and functionality of the parts. The materials that offer advantages in certain applications may not be suitable for others. The results are also influenced by the expertise of the operators. The effect of some of these factors is known and that of others is not.

The results are also influenced by the expertise of the operators. The effect of some of these factors is known and that of others is not. Figure 1 shows a few of the factors which affect part characteristics.First time buyers and users often do not understand which system suits their requirements best. Even machine operators usually possess knowledge about one process only. How can a first time user, or a person with limited experience, go about choosing the process that is best suited for his/her requirements? To answer the many questions a potential buyer/user faces, the obvious course of action would be to compare the various processes by studying data on their performance. However, there are no benchmarking standards in the RP industry at the present time.This paper will address the establishing ofsystematic benchmarking procedures which may be used to generate useful data on some important material and process characteristics ofRP systems.

Qualitative issues such as appearance, form and feel will be important to users with 'touch feel' applications such as design reviews and marketing campaigns. The test part designed to study these will be described here. Some observations about test parts built with four RP technologies, namely, stereolithography (SL), selective laser sintering (SLS), laminated object manufacturing (LOM) and fused deposition modeling (FDM), will be presented.investment casting and functional testing, for example, the dimensional accuracy, build defects such as curl and warpage, and the strength of the parts will be important. Test parts have been designed for measuring repeatability, curl, warpage, creep, shrinkage and tensile strength.

Test parts have been designed for measuring repeatability, curl, warpage, creep, shrinkage and tensile strength. Detailed measurement plans for collecting useful data are still being developed; the design and results of these experiments will be presented at a later date.It is important to stress that both the qualitative and quantitative results will indicate the combined capability of each process and the corresponding material, together with the skill of the operator(s). The long term goal of this project is to be able to understand the effects of process_related, materialJelated and operator_related factors and develop a methodology to study them separately.1462. Previous Benchmarking StudiesAn extensive survey of literature, combined with information from manufacturers, users and service bureaus turned up three benchmarking studies whose results have been published. Only two companies, 3D Systems Inc. and DTM Corporation, publish/furnish information on their test parts.

Only two companies, 3D Systems Inc. and DTM Corporation, publish/furnish information on their test parts. These companies use completely different parts for measuring dimensional accuracy. 3D Systems is the only company to have designed tests for studying different part and process characteristics.Chrysler's Jeep and Truck Engineering conducted a benchmark study [2] in order to decide which system to invest in. A finely detailed speedometer adaptor 1.5"x1.5"x3" in size, shown in Figure 2, was built on six different systems; 3D Systems' SLA 250 and SLA 500, DTM Corporation's SLS 2000, Helisys' LOM 1015, Cubital's Solider 5000 and Stratasys' 3D Modeller. Since Chrysler owned only the SLA 250, they sent the part's STL file to all the manufacturers. For Chrysler, system speed and cost were the most important factors and were studied in detail. Cost was detailed under material, operation, pre and post processing, depreciation and maintenance.

For Chrysler, system speed and cost were the most important factors and were studied in detail. Cost was detailed under material, operation, pre and post processing, depreciation and maintenance. The time study included pre-processing, build and post-processing time on each system. Accuracy was not studied in detail, parts were simply measured to ensure that they were within specifications. Nor did Chrysler study the strength or surface finish of the parts.In 1994, Chrysler sent the same part to RP manufacturers worldwide [4]. The formula for cost comparison was updated but the objective to assist users or purchasers of RP equipment in selecting the fastest and/or cheapest system remained the same. The study concluded that there are It is recommended that each company no set rules to choose anyone RP system over others.

The study concluded that there are It is recommended that each company no set rules to choose anyone RP system over others. review factors like end use of models, urgency of turn around time, availability of capital and trained personnel and location of equipment before selecting any technology.Aubin [3] documents a study of RP technologies from all over the world. The report lists system, maintenance and training costs, build volumes and expected accuracies. Technical capabilities of available systems were characterized and technologies under development were reviewed. A benchmark study was carried out using the Intelligent Manufacturing Systems (IMS) part shown in Figure 3. 4 parts were sought from each company, fabricated from an STL file. The participants were allowed to build a maximum of 6 parts and present the best 4, indicating how many they had built. However, information on the measurements obtained from the parts is proprietary and is not available.

However, information on the measurements obtained from the parts is proprietary and is not available. The comparison of the pre-processing, building and post processing times indicates that the processes studied compare differently in these areas.Van Putte [1] describes a study conducted by Eastman Kodak to study the capability of five RP processes to faithfully reproduce features on a test part. The part, shown in Figure 4, was originally designed to see how various computer-aided design (CAD) / computer-aided manufacturing (CAM) packages could design, alter, analyze and machine typical Kodak components.Part drawings were made on Pro-Engineer and Aries. The STL files were sent to the service bureaus of five companies: 3D Systems, DTM, Cubital, DuPont and Helisys. The parts were measured using a coordinate measuring machine (CMM), vernier calipers and an optical comparitor.

The parts were measured using a coordinate measuring machine (CMM), vernier calipers and an optical comparitor. Sixteen X-Y and four Z measurements were taken on each part.Subtracting the CAD dimensions from the measured ones gave a set of data for each part. Using a statistical software package, the data was analyzed in 2 sets, one for the X-Y plane and the other for the Z-plane, to determine the capability of each process to replicate the original CAD model. A shrinkage correction factor was applied by changing the limits for the capability analysis in such a way as to maximize the performance of each part. A "'WAR-PAGE' chart designed by Cubital was used to subjectively assign a warpage value to the base of each part by comparing it to the lines on the chart.Van Putte himself voices many concerns about the study. Only one part was built by each process.

Only one part was built by each process. The test part design consisted of features important to Kodak and may have been more favorable to some processes than to others. The same software was not used to generate part drawings. All these factors limit the usefulness of the results to other users who are looking for benchmarking data.147While the Chrysler and Kodak studies are interesting, they are primarily of use only to themselves and the data is of little use to other users of these RP systems whose typical parts may be significantly different, as also the factors critical to their applications. The objective of the present study is to fill this gap by designing experiments which can be used to systematically compare RP systems.~ ~roup .of European companies has also conducted a benchmark s!udy [8]. A test part containIng thIn walls, tapers, slots, holes and free form surfaces was desIgned and built by 12 vendors.

A test part containIng thIn walls, tapers, slots, holes and free form surfaces was desIgned and built by 12 vendors. The results are available but have not been reviewed by the authors at this time.3. Qualitative IssuesA simple test part (shown in Figure 5) which contains only cylinders, cones and prismatic boxes, was designed. Each part has four cylinders, tilted at 0, 30, 60 and 90 degrees from the vertical axis. These were used to study the effect of tilting features. A stepped cone with four sections of different cone angles was used to study stair-stepping. The prismatic boxes were used to study straightness and parallelism of edges and warpage of flat surfaces.Parts were built with 6 materials on the 4 systems mentioned in section 1. A study of the parts and processes, combined with technical discussions with machine operators [5,6,7], provided an insight into various pre-processing, building and post-processing issues.

A study of the parts and processes, combined with technical discussions with machine operators [5,6,7], provided an insight into various pre-processing, building and post-processing issues. Such understanding is essential in developing benchmarking standards.3.1 Pre-Processing IssuesPart building is affected by the inherent errors in converting data from a CAD file to a format acceptable by an RP system. Any benchmarking process will have to ensure that the same CAD software and file format are used. However, this still does not rule out differences in interpretation of the accepted format by the system software.Build layouts for this study were created as assemblies on Aries by using multiple copies of the test part. This created a concatenation of STL files for each assembly, which the SLS and LOM software found hard to read [6].

This created a concatenation of STL files for each assembly, which the SLS and LOM software found hard to read [6]. Also, in one instance, the FDM software read one cylinder as being separate from the rest of the test part, in spite of the cylinder being 'unioned' to the base of the part [6]. Subsequently, the part built without problems.An important finding is that while the STL format was accepted by all the systems used in the study, the method of constructing the drawing file on a CAD system was found to affect part building on the LOM process. The LOM process is different from the other processes in that it builds parts by cutting away excess material while the others build parts by adding or fusing material where required. The software on the LOM machine recognizes different entities in an STL file and cuts them separately from the sheets of paper.

The software on the LOM machine recognizes different entities in an STL file and cuts them separately from the sheets of paper. Therefore, on the test part, the cylinders were separated from the base since they were drawn separately and 'unioned' into the base. This would not have happened if the drawing had been constructed by subtracting geometry from a block; the test part would then have been treated as one entity [6,7].Constructing the part drawing with the LOM process in mind will solve this problem but we can only speculate whether the laser on the SL and SLS machines will cure parts differently, based on whether two features were drawn as separate entities or constructed out of one block. If some areas are cured/sintered twice over, parts may be affected in ways not yet known.3.2 Part Building IssuesAesthetics: The SL parts (5154, Weave) looked the best, except for 'swirls' on some inclined cylinders.

The SLS (polycarbonate), FDM (wax), SLS (nylon), SL (5180, Quickcast) and LOM parts rank in order on the basis of appearance. Laser scoring and cross-hatching are causes for the LOM parts to rank low.Geometric features: All 4 processes appear capable of building cylinders at different angles. The results can be expected to be different if hollow cylinders are built, especially with thin walls.148Stair_stepping was not a problem on any of the cones on any machine. However, only the SLA built all cone tips well. All other processes built some parts where the sharp point was not achieved. Edges were sharpest on LOM parts, followed by SL (5154, Weave), SL (5180, Quickcast), SLS (nylon), FDM (wax) and SLS (polycarbonate) parts.Material issues: Problems exist with materials on all 4 systems. Building nylon parts on the SLS process requires very tight temperature control [6].

Building nylon parts on the SLS process requires very tight temperature control [6]. Temperatures even slightly higher than the recommended value harden the loose nylon powder which becomes difficult to remove, affecting appearance and accuracy. In some areas on the test parts, the nylon was so hard that it could not be removed even by sand blasting. A smaller build volume is also recommended for nylon parts [6].With the LOM process, delamination occurs between layers on some parts. Variation in paperquality, specifically coating inconsistency, is suspected to be the one of the causes for this [7].The new resin (5180), used on the SLA, has higher surface tension than the old ones. Layers It also increases build timefiner than 0.006" are hard to build in styles other than Quickcast. because there is a 'pre-dip' delay between layers while the previous layer cures.While building nylon parts on the FDM process, the nozzle leaves fine "fuzzes" on parts.

because there is a 'pre-dip' delay between layers while the previous layer cures.While building nylon parts on the FDM process, the nozzle leaves fine "fuzzes" on parts. This can be controlled by using recommended nozzle sizes and velocities, but the problem is still more evident than with investment casting wax [5].Operation issues: While all four technologies permit build interruption, none of them continue a build successfully after a long delay. Unexpected interruptions occurred on the LOM and SLS machines. A paper jam stopped the build on the former while exhaustion of the nitrogen supply from the supply cylinder was the cause on the latter. An improved paper transport mechanism will help reduce operator monitoring for the LOM process.

An improved paper transport mechanism will help reduce operator monitoring for the LOM process. For large builds on the SLS process, a continuous supply of nitrogen from a plant, as opposed to using a cylinder, may become necessary.Build time: Although the present study did not involve comparison of build times, some factors which limit process speed are highlighted for interest. The FDM process took the longest time for building a set of parts. Faster builds can be achieved by increasing the space between successive nozzle paths, which determines how densely the material is laid out. However, the mechanical traverse of the nozzle is always slower than the scanning speed of a laser using a set of mirrors. On the LOM process, even though a laser is used, it traverses mechanically, again limiting build speed. Also, since excess material has to be removed after the build is complete, cross-hatching has to be carried out over the entire block.

Also, since excess material has to be removed after the build is complete, cross-hatching has to be carried out over the entire block. For small parts, a significant portion of build time will be spent on this.Build failures: In the first set of parts built on the SLS machine with polycarbonate, many of the inclined cylinders and some of the horizontal cylinders broke loose. Some cylinders were also broken into dissimilar sections, appearing to indicate that the sintering had failed to hold material together in those sections. In a subsequent build, parts built with the same material were intact. The cause for the failure is not known.The first build with the Quickcast style (resin 5180) also failed unexplainably. Fine strands of material were found all over the parts, completely covering them with 'fuzz'. All features were stretched in the Z direction; the parts were 1.5 to 2 times taller than their actual size.

All features were stretched in the Z direction; the parts were 1.5 to 2 times taller than their actual size. The next build was good, though no paranleters were changed.3.3 Post-Processing IssuesAll RP processes have recommended procedures which are followed before part building is truly complete. In addition, operators develop their own skills and techniques from experience. To some extent, post-processing is dictated by the end use of the part. Parts for design reviews may be smoothed and painted, Quickcast parts meant for investment casting are carefully drained of uncured resin, and so on. Operator skills influence final part accuracy and finish to varying degrees on different systems. This influence is highest on the LOM parts.Removing excess material from the LOM parts is difficult where geometries are intricate. Special tools are not available. Operators use suitable wood working tools.

Special tools are not available. Operators use suitable wood working tools. Not knowing the149exact geometry can make cleanup even harder, with the danger of accidentally breaking part features being very high for delicate sections. When features are broken, they can be glued back on with wood glue but the effect of this glue on dimensions, especially in the Z direction, needs to be studied. The process is best suited for large, simple parts.FDM parts require little or no processing. Fine strands of material, which may appear on some parts, can be blown away with an ordinary hair dryer [5]. SLS parts rank second in ease of post_processing, the most difficulty is encountered with cleaning hardened nylon. SLA parts usually require more time than FDM and SLS parts since post_processing involves support removal, draining of uncured resin, curing in an oven and cleaning with chemicals.4. Quantitative IssuesThe factor of greatest interest to this study is repeatability.

Quantitative IssuesThe factor of greatest interest to this study is repeatability. Repeatability is really a machine issue; two supposedly identical machines, though theoretically of the same capabilities, will still It is important for users to understand how consistently a particular have their differences. machine will build a part when built at different times, platform locations, orientations, by different operators using different materials. It is reasonable to expect variations as mechanical parts wear on the machines. By following a standardized test procedure, it is possible to monitor a machine's capabilities and make adjustments wherever possible to obtain builds that are known to be within the machine's capabilities.

A benchmarking procedure will be developed in this light using the test part which was used for the qualitative study described in section 3.Other factors of interest include build defects such as curl and warpage, material strength(tensile), and part deterioration over time (creep, shrinkage).Warpage is defined as the out of plane deviation of flat surfaces. The test part is, once again, the one described in section 3. A CMM can be used to measure the coordinates at many points on the top surface of each prismatic box. By fitting a plane through these points, the deviations found at the inspected sections can be used to calculate warpage.For determining tensile strength, it is proposed that ASTM standard 638-72 flat tensile testbars be built and tested.Curling is the deformation that occurs in parts with geometries such as cantilever lengths and is due to internal stresses which develop during part building.

Creep is defined as the dimensional change that occurs over time due to molecular changes and gravity, that is, the weight of the part. Two types of shrinkage can be studied, linear and bulk shrinkage. Linear shrinkage is defined as the change in part dimensions over time. Bulk shrinkage is defined as the change in volume of a part over time. Test parts and procedures to verify performance of RP systems on these factors will be developed.AcknowledgementsFord Motor Co. is thanked for financial support, for the support and cooperation of many staff members in the rapid prototyping laboratory at the Alpha Advanced Manufacturing Division in building all the test parts and for the cooperation extended during one of the authors' (Jayaram's) Persall at the Product Realization Laboratory, visit to their facility. The authors also thank Ms. Clemson University, for all her suggestions and help in our 'part cleaning workshop'.References1.

The authors also thank Ms. Clemson University, for all her suggestions and help in our 'part cleaning workshop'.References1. Van Putte, D.A., "A Brief Benchmarking Study of Rapid Prototyping Processes", Proceedings of the Third International Conference on Rapid Prototyping, Dayton, OR, June 1992,pp. 251 263.2. Wohlers, T., "Chrysler Compares Rapid Prototyping Systems", Computer-Aided Engineering, October 1992, pp. 84-90.3. Aubin, R.F., "A World Wide Assessment of Rapid Prototyping Technologies", Report No. 94-13, January 1994, United Technologies Research Center, 411, Silver Lane, MS 129-48,150East Hartford, CT 06108.4. Schmidt, L.D., "A Benchmarking Comparison of Commercial Techniques in Rapid Prototyping", Proceedings of the Fifth International Conference on Rapid Prototyping, Dayton, OH, June 1994, pp. 333-351.5. Abramson, R., Alpha Advanced Manufacturing Division, Ford Motor Co., Detroit, MI., Personal Communication, April-July 1994.6.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperBenchmarking the Tensile Properties of Polylactic Acid (PLA) Recycled Through Fused Granule Fabrication Additive ManufacturingDawood Al Nabhani*, Ali Kassab†, Osama Habbal*, Pravansu Mohanty*, Georges Ayoub†, and Christopher Pannier* *Department of Mechanical Engineering, University of Michigan – Dearborn, MI 48128 †Department of Industrial and Manufacturing Systems, University of Michigan – Dearborn, MI48128AbstractTo progress toward a circular economy of thermoplastic polymers, the adoption of 3Dprinters to make functional articles can facilitate distributed recycling. To this end, the mechanical degradation of polymers through multiple recycling cycles must be quantified.

To this end, the mechanical degradation of polymers through multiple recycling cycles must be quantified. This work presents a procedure and benchmark dataset of tensile property degradation for polylactic acid (PLA) feedstock in multiple recycling passes with a fused granule fabrication process. To establish recycling with minimal processing (shredding and sieving), modifications were required to the granule feeding hopper of the 3D printer. Two distinct orientations were chosen to obtain tensile test coupons. These coupons were die-cut from machined 3D printed rectangular cross-section tubes, with one orientation along the bead (0°) and the other perpendicular to it (90°). Tensile properties are presented for 3D printed virgin material and one, two, three, and four passes of recycling. In terms of print orientation, the results indicate that samples pulled at 0° and 90° exhibited similar mechanical properties.

In terms of print orientation, the results indicate that samples pulled at 0° and 90° exhibited similar mechanical properties. However, there was an average decrease of 3.1% in ultimate tensile strength and a 1.7% decrease in elastic modulus for the samples along 90° orientation for all recycling passes. The samples along 0° demonstrated a 13.7% higher strain at fracture compared to those along 90°. Regarding the number of recycling passes, the findings suggest that the mechanical characteristics of PLA remain largely unaffected even after undergoing four recycling cycles. However, when the material is pulled in the direction of the bead, a 3.09% decrease in ultimate tensile strength is observed in the fourth recycling pass. The elastic modulus and strain at fracture did not exhibit a clear trend.

The elastic modulus and strain at fracture did not exhibit a clear trend. It is important to note that the testing results display some variability, which can be attributed to a combination of stochasticity in the printing process and the preparation procedure employed.Keywords: Additive Manufacturing, Recycling, PLAIntroductionThe concept of circular economy (CE) in thermoplastic polymers has received global attention, including the proposed distributed recycling by additive manufacturing (DRAM) of thermoplastic polymers [1]. In particular, the fused granule fabrication (FGF) process holds great potential for DRAM by utilizing shredded feedstock and eliminating the requirement to produce a spooled filament with high tolerance [2]–[4]. The demand for distributed recycling is growing due to the projected increase in global waste generation. It is estimated to rise from 2.01 billion tons in 2016 to 2.59 billion tons in 2030 and further to 3.40 billion tons by 2050.

It is estimated to rise from 2.01 billion tons in 2016 to 2.59 billion tons in 2030 and further to 3.40 billion tons by 2050. Notably, 12%22of this waste consists of municipal plastic waste [5]. Therefore, open-source 3D printers are proliferating to give access to many users at all levels including individuals, community groups, and businesses to recycle thermoplastic waste at a local level [6],[7]. However, although recycling offers promise for creating a circular economy, it has its own limitations. First, for the purpose of mechanically recycling plastics, the material needs to be melted by significant motion of the chains. Polymers consist of various types of chains, and only thermoplastics, which are made up of linear chains, can be remelted. This limits the range of plastics that can be effectively mechanically recycled. Second, polymers experience degradation over time due to factors such as high temperatures, light exposure, moisture, humidity, or microorganisms.

Second, polymers experience degradation over time due to factors such as high temperatures, light exposure, moisture, humidity, or microorganisms. This degradation causes the material to lose its properties, reducing the likelihood of being recycled and meeting the necessary quality standards expected by the recycling community [8]. Polylactic acid (PLA) is a biodegradable semicrystalline polymer derived from biomass [9]. PLA has been widely used in various applications, particularly in consumer products, compostable products, and 3D printing filament, due to its favorable characteristics such as a low glass transition temperature and minimal shrinkage during 3D printing. Therefore, recyclability of PLA is of high interest. Thus, it is important to consider the mechanical properties of PLA, as they may undergo degradation after multiple recycling cycles [10].

Therefore, recyclability of PLA is of high interest. Thus, it is important to consider the mechanical properties of PLA, as they may undergo degradation after multiple recycling cycles [10]. Understanding the mechanical performance is crucial in determining the number of recycling passes before the material's mechanical integrity is compromised and it becomes unsuitable for producing products.Although there is a rising interest in the recyclability of thermoplastic polymers in the field of 3D printing, prior investigations have predominantly prioritized mechanical recycling methodologies, with a particular emphasis on injection molding techniques [11]. Several studies explored the behavior and properties of 3D printed parts after multiple recycling passes. Cieslik et al. examined the effects of multiple reprocessing cycles on PLA with a conductive composite.

Cieslik et al. examined the effects of multiple reprocessing cycles on PLA with a conductive composite. They investigated variations in extrusion temperature and the number of recycling passes, observing a decline in thermal stability and material strength. During this investigation, instead of utilizing a 3D printer to extrude the filament, the material was broken down and melted at temperatures similar to those used for the material printing [12]. In their study, Tanney et al. aimed to evaluate how multiple extrusion cycles affect the properties of 3D printed parts and explore the feasibility of using a combination of new and recycled materials to maintain material integrity. To simulate the typical recycling process, the authors followed a three-stage approach. First, they pelletized the filament by feeding it into a chamber with a wood routing bit controlled by a drill press. Then, they recycled the filament by using a Filabot Original filament extruder to create a new filament.

Then, they recycled the filament by using a Filabot Original filament extruder to create a new filament. Finally, the recycled filament was used in the printing phase to create the 3D printed parts. The findings revealed a significant reduction in the strength, maximum strain, and toughness properties of the material due to the repeated heat cycles [13]. Anderson conducted an experiment using commercially available PLA and a basic 3D printer. They created and tested tensile and shear specimens to assess their ultimate tensile strength, elasticity, and hardness. The specimens were then recycled into filament using a Filabot EX2 and subjected to a second round of testing. The results showed that 3D printing with recycled PLA using the Filabot EX2 is a viable option. The recycled filament exhibited a slight decrease in tensile strength, an increase in shear strength, and a decrease in hardness. The tensile modulus of elasticity remained unchanged.

The recycled filament exhibited a slight decrease in tensile strength, an increase in shear strength, and a decrease in hardness. The tensile modulus of elasticity remained unchanged. However, the recycled filament showed more variability in the results compared to the virgin specimens [14].23An important aspect in 3D printing is anisotropic mechanical properties in a printed part. This is a well-established phenomenon often considered a drawback in additive manufacturing, especially in the fused filament fabrication (FFF) and fused granule fabrication (FGF) processes. Gradinaru et al. [15] found that PLA mechanical properties differed along the bead and perpendicular to it. However, the magnitude of the difference varies across different studies examining similar properties. Although there is an agreement that 3D printing materials exhibit direction-dependent mechanical properties, the specific results vary due to differences in sample preparation and the methodologies followed.

For instance, in Alexandre et al.'s research [16], a substantial difference in tensile strength was observed between the 0° and 90° orientations, with the 0° orientation showing five times the strength of the 90° orientation. In another study, conducted by Grant et al. [17], the tensile strength was found to be twice as high in the 0° orientation compared to the 90° orientation. These works show that magnitude of anisotropy varies with printing and testing procedure.The primary objective of this study is to quantitatively assess the degradation of PLAmaterial after multiple recycling passes using FGF additive manufacturing process. By establishing a benchmark dataset that evaluates the degradation of tensile property degradation, this research aims to understand the PLA mechanisms of degradation as it undergoes additive manufacturing recycling across four iterations.

To mitigate potential printing challenges arising from the use of recycled material, modifications were made to the feeding hopper of the GigabotX 2 XLT printer [18] incorporating an upgraded fluidized bed system to minimize occurrence of clogs while printing of the recycled passes.Methodology2.1 Material PreparationIn this study, LX175 virgin PLA pellets were sourced from a single supplier (Filabot) toensure a consistent and standardized material with uniform properties. The pellets have an average diameter of 4.2 mm as shown in Figure 1 with a melting temperature range of 180-210° C [19].

The pellets have an average diameter of 4.2 mm as shown in Figure 1 with a melting temperature range of 180-210° C [19]. To maintain consistency across the recycling passes, the granules were shredded repeatedly to achieve the same average diameter as the virgin pellets with 95% confidence level.ABFigure 1: Determining the Size of the Virgin Pellets(A) and the Recycled Granules (B).24To investigate the mechanical degradation caused by multiple recycling passes, the virginPLA was subjected to a series of recycling passes. Figure 2 presents an overview of the methodology used to produce the dog-bone specimens intended for testing purposes.2.2 Recycling ProcessThe recycling process involved a series of steps, beginning with shredding, sieving,followed by drying, 3D printing, machining and finally punching along two different material directions (0°and 90°).

These steps were designed to mimic real-world controlled recycling condition, enabling the process to be easily scaled for larger quantities of recycled material.Figure 2: PLA Recycling Process2.2.1 ShreddingThe shredding process consisted of two stages. During the first one, the tubes were cut into 3 x 3 inches and fed into the shredder, resulting in granules sizes of approximately half an inch. In the second stage, the shredded PLA granules were subjected to multiple rounds of shredding, a process where they were repeatedly broken down into smaller pieces. This iterative shredding was carried out to continuously reduce the size of the PLA granules, making them more suitable for the printing process. The repeated shredding helped achieve the desired particle size for 3D printing.

The repeated shredding helped achieve the desired particle size for 3D printing. To maintain a contamination-free process, shredder was cleaned between each cycles using compressed air and hand tools.2.2.2 Sievingm m 6 >m m 6 - 5m m 4 - 3m m 3 - 2m m 2 <Discardedm m 33mmUsed for the ProcessThe Sieving process is an essential step for granulesprocess. The purpose of this step is to segregate the shredded granules based on the size and use granules within the size ofFigure 3 : Sieving System253-4 mm as a feedstock to the 3D printer. To accomplish this requirement, a total of five sieves with different mesh sizes were utilized, classifying the granules into the following categories: >6mm ,5-6mm ,3-4mm, 2-3mm, <2mm. The arrangement and sequence of the sieving process can be observed in Figure 3. Any granules that exceeded the desired size were returned to the shredding system until they successfully pass through the 3-4mm sieve.

Any granules that exceeded the desired size were returned to the shredding system until they successfully pass through the 3-4mm sieve. Conversely, granules with a diameter smaller than 2mm were discarded.2.2.3 DryingThe PLA material was dried by a conventional oven at a temperature of 45°C for aduration of 6 hours. This controlled drying process aimed to eliminate any moisture content in both the pellets and granules, thus ensuring optimal printing performance. Moreover, to enhance the preservation of the PLA material and minimize moisture absorption prior to printing, an additional precautionary measure was implemented by storing the material in a vacuum chamber.2.2.4 3D PrintingA Gigabot X 2 XLT 3D printer was used to print the production square tubes represented by the CAD model in Figure 4.The tube was printed with the dimension of 80X80X100 mm (LxWxH) and a thickness of 3mm.

This specific design was chosen to maximize the available wall surface area enabling the cutting of dog-bone shapes for tensile testing. Depending on the material direction along which the sample was extracted, whether 0° or 90°, each wall of the tube could yield approximately 5-6 dog-bone specimens.Figure 4: 3D Printing Tubes2.2.5 MachiningAfter the tubes were 3D printed, they were cut intofour walls, followed by machining using an endmill from both sides to achieve a uniform thickness of 1.5mm, with tolerance of 1.47±0.09 mm. It is crucial to emphasize that ensuring flatness in the machined surfaces is of paramount importance. Any deviations in thickness can potentially affect the outcomes of the tensile tests. Therefore, strict control and thorough measurement of the flatness of all machined surfaces' thickness was carried out during this process. To provide visual insight, Figure 5 presents a cross-sectional view of a 3D printed tube wall, highlighting the bead patterns.

To provide visual insight, Figure 5 presents a cross-sectional view of a 3D printed tube wall, highlighting the bead patterns. The figure serves to illustrate the necessity of machining the tube walls to achieve the desired flatness and uniform thickness as well as to eliminate uneven edges, which might affect testing, as outlined in the research conducted by Cole et al.[15].Figure 5: Cross-sectional View of a 3D Printed Tube Wall2.3 Fluidized Bed for the Feeding SystemIn order to accommodate the 3D printing of recycled PLA granules, a modification wasmade to the feeding system of the GigabotX 2 XLT printer [18]. This modification was essential to ensure proper flow and uniform distribution of the recycled PLA granules during the additive manufacturing process. Figure 6 shows a detailed schematic of the fluidized bed system that was26designed specifically for the Gigabot X 2 XTL.

Figure 6 shows a detailed schematic of the fluidized bed system that was26designed specifically for the Gigabot X 2 XTL. The development of this upgrade was prompted by challenges encountered with clogging due to the irregular shape of the granules, which was expected in the recycling process. After numerous attempted alterations, the proposed modification demonstrated the highest success rate for achieving successful prints with granules. The system consists of a compressed air system delivering air at a constant rate of 3CFM and 115psi (gauge pressure), with the inlet controlled by a motorized ball valve that opens and closes based on the cycle showed in figure 7.

Additionally, Figure 8 shows an image showcasing the installed fluidized bed system and presents the modified CAD design of the fluidized bed.HopperControlCircuitPower SupplyFeed ThroatAir SupplyExtruderFlow Meter 3 CFM 115 PSIAirElectrical WireFigure 6: Fluidized Bed SchematicOpenClosed5.06.1 6.37.417.4 23.5 29.8 37.2 50.0Figure 7: Ball Valve Opening and Closing Cycle27Figure 8: Photograph of the installed fluidized bed feed throat (A). Feed throat CAD design with section view (B).2.4 Tensile Test Coupon PreparationTo ensure precise and accurate results, a procedure was developed to control thepunching process. Figure 9 shows the septs used to punish the dog bone test specimens. The process involved using a manual press machine to punch out dog samples used for tensile testing. To facilitate the punching process, the 1.5mm sheet were heated in a convection oven at 50°C for 60 seconds.

To facilitate the punching process, the 1.5mm sheet were heated in a convection oven at 50°C for 60 seconds. It is important to highlight the criticality of aligning the punching direction accurately. Two specific punching direction were used: 0° and 90° degrees, as clearly illustrated in Figure 2. The sheet was aligned either at 0°or 90° degrees from the reference bead, followed by the drawing of a precise straight line to guide the alignment of the dye cutter.A non-standard dog-bone cutting die used was with a radius shape in the gauge area wasselected. This test samples have a thinner middle section of the samples, increasing the likelihood of failure occurring at the center. A visual representation of the shape and geometry of the cutting die is provided in Figure 10.The tensile properties of the 3D printed samples were assessed using the Instron E1000mechanical testing system, which was equipped with a 1 KN load cell.

Tensile tests were carried out at room temperature and using an engineering strain rate ofs-1.To obtain the engineering stress-strain curves, the engineering stress was calculated by−410measuring the cross-sectional area of the specimen at its narrowest point, which is at the vertical center. The displacement was recorded from the software at the crosshead movement, and the engineering strain was determined using a gauge length of 43 mm, as shown in Figure 10. Figure 11 illustrates the Engineering stress-strain curve of the virgin sample for both the 0° and 90° orientation.28Figure 9: Punching set-up used for the experiment.10167543 (Gauge Length)4R78.540°90°Figure10: Sample geometry and orientation, dimensions in mm.

The gap between the lines in both orientations represents the bead thickness, which is equivalent to 1.5mm.2.5 Additive Manufacturing/3D Printing ProcessThe Gigabot X 2 XTL printer [9] was used for the fused granule fabrication (FGF) process, which was utilized to produce the tubes. Both the virgin PLA material and the recycled PLA granules from each recycling pass were loaded separately into the granule feeding hopper of the printer. For slicing Simplify3D software was used, and the specific printing parameters are outlined in Table 1. Notably, fan cooling was not utilized during the printing process, and the maximum extrude cross-sectional area was optimized, increased from 7 to 12 mm2.29Table 1: Printing ParametersResults and Discussion 3.1 Tensile Testing Method and ResultsIn the context of 3D printing, the layer-by-layer building process often leads to rounded edges and unfilled rectangular cross-sectional areas, resulting in a stepped surface appearance.

To address this, machining techniques are utilized to eliminate these edges and achieve a smooth, flat surface for the final specimen geometry. Without machining, these imperfections may have been overlooked during the calculation of the cross-sectional area for tensile testing. Tensile testing was conducted on both virgin and recycled PLA samples, extracted along two material orientations. Our findings indicate that the print orientation does have a significant effect on yield strength when the specimens are subjected to tensile testing to failure. Our research specifically focuses on examining the adhesion and alignment of bead centers. Within our experimental design, we observed an average reduction of 3.12% in ultimate tensile strength and a 1.66% decrease in elastic modulus for the parts printed at a 90 degrees orientation.

Within our experimental design, we observed an average reduction of 3.12% in ultimate tensile strength and a 1.66% decrease in elastic modulus for the parts printed at a 90 degrees orientation. It is important to note that this decrease was not consistently observed across all recycling passes, which can be seen in the virgin and fourth recycled pass.303.2 Property Degradation70) a p M(6050s s e r t S g n i r e e n i g n E4030200°90°10000.0050.010.0150.020.0250.03Engineering Strain(a)70) a p M( s s e r t S g n i r e e n i g n E60504030200°90°10000.0050.010.0150.020.0250.03Engineering Strain(b) Figure 9: Stress Strain Curve corresponding to the Virgin (a) and 4th Recycling Pass (b) Prints for 0° and 90° OrientationsThe recycling processes involved one, two, three, and four passes.

When comparing different print orientations, there was no evident pattern observed in the ultimate tensile strength or elastic modulus. However, it was consistently observed that samples with a 90° orientation tended to fracture earlier in comparison to samples with a 0° orientation. The samples printed at 90° demonstrated an average strain at fracture of 2.145%, whereas the strain at fracture for the 0° samples was 2.438%, indicating a 13.7% higher strain at fracture.(a)(b)(c)Figure 13: Regression Line over Ultimate Tensile Strength (a), Elastic Modulus (b) and Total Strain at Fracture (c) at 0° and 90°33Referring to Figure 13(a), the samples tested in a 0° orientation exhibited a slight decrease in ultimate tensile strength with respect to recycling pass number. The average ultimate tensile strength of the virgin samples was 60.92 ± 0.87 MPa (2 standard deviation error), which decreased to 59.03 ± 1.08 MPa after undergoing four recycling passes.

The average ultimate tensile strength of the virgin samples was 60.92 ± 0.87 MPa (2 standard deviation error), which decreased to 59.03 ± 1.08 MPa after undergoing four recycling passes. This represents a 3% reduction in the ultimate strength. There is no definite trend in the results shown in Figure 13 (a, b and c) where the correlation coefficient (R2) ranged from 0.02 to 0.12. These values suggest a very weak correlation within the dataset, contrary to prior research that generally observed significant differences in ultimate tensile strength between the number of recycling passes. This outcome is unexpected, and the variation from previous literature may be attributed to differences in the method used for sample preparation in our research.ConclusionThis paper has presented a comprehensive study on benchmarking the tensile properties of polylactic acid (PLA) recycled through the fused granule fabrication additive manufacturing process.

The research established a procedure and provided a benchmark dataset for evaluating the mechanical degradation of PLA through multiple recycling passes.The findings indicate that there is no notable degradation in the mechanical properties ofthe tensile tested specimens throughout four recycling passes. There is no visible trend evident regarding the elastic modulus or strain at fracture, displaying neither an increase nor a decrease, which was surprising since the expected results should have shown a significant difference in different orientations. However, a decrease of 3% in ultimate tensile strength was observed solely in the case of the 0° orientation. Although the data suggest potential trends, the available experimental data are insufficient to confirm them definitively.

Although the data suggest potential trends, the available experimental data are insufficient to confirm them definitively. Therefore, further investigation is necessary, which could involve conducting additional tests with a larger sample size or considering modifications to the experimental protocol.The sample geometry of stacked individual beads with bead edges eliminated by machining allowed us to examine the layer-to-layer adhesion and alignment of bead centers in our tests. It is unknown whether our finding that samples pulled in the 0° and 90° orientations displayed comparable mechanical properties would extend to samples including lateral bead edges with the same processing conditions.

Many FGF applications use structures with multiple perimeters and nonzero infill, unlike our samples composed of single bead perimeters with no infill, and thus additional testing would be needed to understand mechanical properties of those structures.In comparing the different recycling passes, we observed that there were no distinct patterns indicating significant degradation. This suggests that the recycling, printing, and milling processes employed in our study do not lead to substantial degradation, even after four recycling passes. These findings indicate that the PLA material used in our study exhibits high recyclability. It is important to note that the controlled experimental conditions we used differ from the real-life conditions that parts may encounter during their usage.

It is important to note that the controlled experimental conditions we used differ from the real-life conditions that parts may encounter during their usage. Factors such as moisture, UV radiation, chemical attack, bacteria, abrasion, dirt/particles, or contamination with other polymers during recycling can affect the performance and properties of recycled materials. Our study focuses on the optimal scenario for the circular economy, demonstrating that thermoplastics, specifically PLA, can withstand four recycling passes without encountering34significant issues. However, we did observe slight variations in properties from one recycling pass to another. This highlights the stochastic nature of thermal processing and the influence of material history. In our study, the extrusion temperature and duration were carefully controlled to ensure the melting of crystals and restoration of the microstructure to an amorphous state (normalized).

In our study, the extrusion temperature and duration were carefully controlled to ensure the melting of crystals and restoration of the microstructure to an amorphous state (normalized). Additionally, the passive cooling in our specific part geometry likely prevented significant crystallization. While our research showcases the potential of PLA for multiple recycling passes under controlled conditions, it is important to consider the real-world factors and challenges that may impact the performance and properties of recycled materials. Further studies are needed to explore the impact of environmental factors and real-life usage conditions on the recyclability and performance of thermoplastics.The benchmark dataset presented in this study provides valuable information regarding the mechanical degradation of PLA through fused granule fabrication additive manufacturing and multiple recycling cycles.

vel Cutting Methods and Cutting Trajectory Control for Steel Laminations Used in Toolingby Daniel F. Walczyk Department of Mechanical Engineering, Aeronautical Engineering and Mechanics Rensselaer PolYtechIlic Institute Troy, New YorkAbstractBevel cutting of steel laminations used in profiled-edge laminated tooling allows for a more accurate representation of the intended die surface, since stair-stepping at the edges is eliminated. Based on experiments involving maximum cutting speed, bevel angle and kerf quality, the three recommended methods for bevel cutting steel are (best to worst) pulsed Nd:YAG laser with hard-optic delivery, abrasive water jet, and machining with the flute edge of an endmill. For each method, bevel angles of up to 80 degrees are possible. Further experimentation was used to determine the optimal process parameters for kerf quality, with constant cutting speed being one of the main requirements.

Further experimentation was used to determine the optimal process parameters for kerf quality, with constant cutting speed being one of the main requirements. Finally, a new techIlique to assure constant velocity along the entire lamination cutting trajectory is developed.1. Introduction and Background Alllamination-based RP methods begin with intersecting a finite number of specified parallel planes (usually equally-spaced) with a 3-D solid or surface model. The resulting collection of contour or profile curves define how the laminations are to be machined and assembled. When laminations are cut perpendicular to their surface (i.e., zero-order approximation), the assembled array of laminations has a stair-stepped surface as shown in Figure 1(a). Thin laminations (e.g., 0.05 mm) that are common with some RP methods like Laminated Object manufacturing (LaM), only require surface finishing to remedy this problem.

Thin laminations (e.g., 0.05 mm) that are common with some RP methods like Laminated Object manufacturing (LaM), only require surface finishing to remedy this problem. With larger models, thicker layers (e.g., 10 mm) are typically used and the stair stepping problem becomes more troublesome. Common remedies in this latter case are to perform secondary machining on the stair-stepped surface or fill in between lamination edges with some hard material (e.g., epoxy).Many have recognized the advantages of cutting lamination edges with a varying bevel instead of simply perpendicular to the surface. Essentially, the lamination edge is defined by connecting two adjacent contour (closed curve) or profile (open curve) lines with a bevel. As shown in Figure 1(b), bevel cutting allows for a first-order or piecewise continuous approximation of the tool or model surface between lamination interfaces.

As shown in Figure 1(b), bevel cutting allows for a first-order or piecewise continuous approximation of the tool or model surface between lamination interfaces. Prior work with thick lamination beveling includes [Beyer et a1.,1988] making laminated models with a laser; [Weaver,1991] making metal tooling with a laser; [Berman, 1991] making models with a routing head; [Glozer et a1.,1993] cutting aluminum with an EDM for plastic injection molds; [Walczyk,1994] cutting steel with a laser, AWJ cutter, and machining head for sheet metal forming dies; [Lee et a1.,1996] cutting polystyrene foam with a electrically heated wire; [Zheng et aI., 1996] making models with a laser; and [Hope et aI., 1997] making models with an AWJ cutter (Styrofoam).As seen in Figure 1(c), bevel cutting of a particular lamination requires line-of-site machining instructions defined by a series of position point and unit directional vector pairs (P), V).

There are several algorithms that have been developed for defining lamination edge bevels and suitable cutting trajectories. [Keppel,1975] developed a heuristic-based algorithm for obtaining an optimal approximation of the bevel surface defined by randomly distributed points along adjacent contour lines (Le., contours are defined by co-planar polylines). The surface approximation is made up of a collection of triangular facets that stitch the adjacent contours together. Based on the [Keppel,1975] graph-search techIlique, [Fuchs et a1.,1977] developed a faster algorithm that does not utilize any heuristics, thereby allowing more options for choosing optimizing criteria. A "minimum square span length" between nodes on adjacent contours was used by [Newman et a1.,1995] as the cost function in an optimal surface reconstruction. Furthermore, [Newman et a1.,1995] used the concept of a cutting contour, defined by the midpoints of the spans, to describe a constant velocity cutting trajectory.

Furthermore, [Newman et a1.,1995] used the concept of a cutting contour, defined by the midpoints of the spans, to describe a constant velocity cutting trajectory. [Zheng et a1.,1996) used a "smoothed" path between two adjacent contours that eliminates the rapid velocity changes corresponding to the transitions between the [Lee et aI.,1996] simply defined the cutting trajectory as the exact geometry of the two computed triangular facets. adjacent contours. Instead of defining a lamination by two adjacent contours, [Hope et a1.,1997] defines a lamination by its mid-contour curve and the associated surface normal, curvature, and tangency data taken directly from the NURBS surface model. This information is used to establish a linear approximation of the local surface121curvature and, hence, the beveled lamination edge. lamination.Furthermore,the user can defme the thickness of each2.

lamination.Furthermore,the user can defme the thickness of each2. Bevel Cutting of Profiled-Edge LaminationsProfiled-Edge Lamination (PEL) tooling (see figure 2(a)) has been shown to have distinct advantages over of automated assembly, registration, securing, and re- EL lamination edges are bevel cut, the issues of developing and dynamically desirable cutting trajectories for common tooling materials (e.g., steel) must be successful implementation of this method in industry. Therefore, the focus of this paper is to compare vanous methods for cutting metal laminations (specifically steel which is arguably the most common tooling material u discuss the associated cutting trajectory requirements, and describe a cutting trajectory algorithm which fulfi s these requirements.contoured lamination tooling (see Fi n:ac.

kme ,2(b)) .tions [Walczy ,1998].Desired Surface~L,m;,,".",Z ~ -·: ~_.(b)laminationnearside profile curve(c)Figure 1 - (a) Stair-stepping of tool surface from perpendicular-cut lamination edges, (b) smoother surface from bevel cutting, and (c) line-of-site machining instructions needed for bevel cutting.Profiled Edge.....PEL Lamination··I""" ..'::..."."Lemlnatlon __RegistrationEdgeTension Rod'/Lemlnatlon Mounting & Clamping Frame(a)(b)Figure 2 - (a) PEL tooling and (b) contoured lamination tooling,3. Bevel Cutting Methods for Steel LaminationsThe most promising methods for bevel cutting steel laminations are machining with the flute-edge of an endmill, abrasive water jet cutting, plasma-arc cutting, and laser cutting [Walczyk,1998]. To determine how rapidly and accurately they will machine steel PEL laminations, each of these methods was investigated through a series of cutting experiments at bevel angles of 0, 30, and 60 degrees.

The basis of evaluation for each beveling method will122be the quality characteristics of the bevel cut as defmed by Figure 3. A narrow kerf wid~h is ~esirabl~ so th~t fi~e details can be cut into the lamination edge. A low surface roughness of the cut surface IS desIrable smce thIS wIll decrease the amount of grinding and polishing time that is necessary for the assembled PEL die forming surface. Minimal or no edge burring on the farside of the kerf is desirable since it must be removed before the die laminations can be assembled into a complete die. Deburring increases the overall fabrication time. The lamination material used for every cutting experiment was 1.5 mm thick SAE 1010 cold drawn steel sheet.ex. Ve max.

Deburring increases the overall fabrication time. The lamination material used for every cutting experiment was 1.5 mm thick SAE 1010 cold drawn steel sheet.ex. Ve max. cutting speed Ra = surface finish (Effects the time required to finish a die)bevel angle (0,30,60 deg.)W = kerf width(Can affect the smallest achievable gcometrical detail)HD = burr or dross height (Burr must be removed prior to die assembly) e = taper angle(Must be compensated for with cutting head orientation)Figure 3 - Quality characteristics of a bevel cut.80 grit abrasive particlesLamination(a)abrasive water jetFigure 4 - Lamination cutting methods including (a) machining with an endmill, (b) abrasive waterjet cutting, and (c) laser cutting.3.1 Machining with the Flute-Edge of an EndmillAs seen in Figure 4(a), a profiled-edge with a compound bevel can be machined using the flute-edge of an endmill mounted in a 5-axis machining center.

According to a series of bevel cutting experiments using a 6.4 mm diameter endmill, beveling the 1.5 mm thick steel laminations with the endmill's flute-edge leaves very good surface finishes but cutting speeds are relatively low. Slower cutting speeds (denoted as Vc) and higher spindle speeds (denoted as N) yield the best surface finish (Ra) and the smallest burr (HD). The maximum Vc and N cutting and spindle speeds for cutting bevels is highly dependent on the material's machinability. Although the presence of coolant did not seem to have any real effect on surface finish and edge burring, it should be used when cutting metal to extend the life of the tool and to minimize the HAZ of the cut. The extent of burring on the climb-milling side of the cut is extensive (see Figure 4a) and it gets longer as the bevel angle increases.

The extent of burring on the climb-milling side of the cut is extensive (see Figure 4a) and it gets longer as the bevel angle increases. Due to the high cutting forces from the large amount of material removed (Le., width of the tool), there is also significant deflection of the lamination during machining which exacerbates any chattering problem and decreases machining accuracy. For comparison purposes, the average surface roughness, burr height, and maximum cutting speed (and spindle speed used) for each of the bevel angles are listed in Table 1.123Bevel Angle (a) Ave. Ra (IJ,m) Ave. HD(mm) Max. Vc (m/min)Table 1 - Experimental resu ts or ute-e Ige en 30° 4.6 2.7 0.46 (at N=3000 rpm)1 flfldm'll'1 mg 0d0° 1.3 0.3 0.46 (at N=3000 rpm)fbeve s.60° 3.1 8.6 0.25 (at N=2000 rpm)3.2 Abrasive Water Jet (AWJ) CuttingAs seen in Figure 4(b), an abrasive water jet (AWJ) is a non-contact method for cutting bevels into die laminations.

The control factors that affect the cut quality are cutting velocity Vc, upstream water pressure P , and size of the abrasive particles. An OMAX JetMachining System was used for the AWJ cutting experiments. The cutting nozzle diameter and standoff distance from the lamination used for all experiments was 0.76 and 1.27 mm, respectively. Since the AWJ manufacturer strongly recommended that an abrasive particle of particular size be used, specifically 80 grit (0.267 mm average particle size), this was not used as a process parameter.It is difficult to generalize about what effect Vc, P w and their interaction has on Ra because the trends vary with every bevel angle. However, Vc and P w have a significant effect on the kerf width w, but in opposite directions. Specifically, W tends to increase with higher PH' and tends to decrease with higher Vc' The effect that Vc and PH' have on HD is inconclusive.

Specifically, W tends to increase with higher PH' and tends to decrease with higher Vc' The effect that Vc and PH' have on HD is inconclusive. Since there appear to be no optimal process parameters for bevel cutting with AWJ in terms of Ra and HD , the author suggests using a maximum Vc and as Iowa Pw as is practical. Average values of various quality characteristics from the AWJ experimental bevel cuts are listed in Table 2.From the experiments, a number of general effects were noticed.AWJ cutting is a very good non-contact method for machining narrow cuts into die laminations. Because it is non-contact and low forces are imposed on cutting area, there is negligible deflection of the lamination during cutting. Aside from a decrease in the maximum Vc, AWJ cutting seems to have no problems cutting high bevel angles. According to [Tikhomirov et al.,1992], the maximum Vc for this process is highly dependent on the yield strength of the material.

According to [Tikhomirov et al.,1992], the maximum Vc for this process is highly dependent on the yield strength of the material. However, the measured maximum Vc for 1.5 mm thick thickness lamination material made of mild steel and much harder tool steel are close in value. The surface finish of the bevel cut is slightly rougher than flute-edge endmilling, but more consistent as process parameters change. The burr left on the far edge of the kerf is much less than CNC-machining with an endmill especially at high bevel angles. However, the kerf made with AWJ has a large taper, around 10° for all the bevel angles tested.

However, the kerf made with AWJ has a large taper, around 10° for all the bevel angles tested. The kerf taper angle may decrease when thicker laminations are cut and/or when the upstream Pw is increased.Even though the cut from an abrasive wat~rjet introduces error in the cut because of the drag line, taper angle of the kerf, and variable kerf width; [Matsui, 1991] has successfully compensated for these errors to increase the precision of the AWJ process as shown in Figure 5. Compensation for these errors involves offsetting the centerline of the kerf by half of it's width which is a function of Vc, rotating the cutting nozzle slightly about the direction of travel (y-axis) by half the taper angle, and tilting the cutting head backwards (about x-axis) to minimize the effect of the lag line.Table 2 - Experimental results for AWJ bevel cutting of 1.5 mm thick steel.a Ave. Ra (/Jm) Ave. w(mm) Ave. HD(mm) effective thickness of material being cut (mm) Max.

Ra (/Jm) Ave. w(mm) Ave. HD(mm) effective thickness of material being cut (mm) Max. Vc (m/min) kerf taper0° 4.8 0.62 0.0611.470.34 10°30° 5.1 0.63 0.1001.700.31 11.5°60° 4.3 0.80 0.0532.950.20 9°124wz" p-4 ~, ))))))~zcuttingcenterline of cuty.... direction~id' ~cut L.I~~ ""-8--:.1 b) taper angle Figure 5 - Compensation methods for AWl cutting error.c) lag linea) kerf width3.3 Plasma-Arc CuttingUsing a plasma-arc is another non-contact cutting method that can be used to cut bevels into steel laminations. The system used for the bevel cutting experiments was a Hypertherm HD-I070 HyDefinition Plasma unit. The control variables in plasma-arc cutting are the plasma gas type (02 or Air), the plasma gas pressure PG, the standoff distance of the cutting nozzle from the lamination HSD' and the cutting velocity Vc' Extensive experimentation was abandoned after it became apparent that a plasma-arc was not suitable for beveled cuts.

Bevel cuts at angles up to 45° yielded very large tapers (> 10°) in the kerf, excessive dross (HD=0.9 mm) that was welded to the metal, and a large heat affected zone (HAZ). The large HAZ caused the cut edge of the lamination to warp slightly. The maximum bevel angle achieved was 60° but the kerf edge closest to the cutting nozzle was consistently obliterated from an over-zealous self-burning l of the metal during cutting.The attractive feature of plasma-arc cutting is the maximum allowable cutting speed which is typically 5 to 10 times faster than that of machining and water-jet cutting. Unfortunately the problems of excessive kerf taper and the tendency of kerf edge nearest to the cutting nozzle to experience self-burning make plasma-arc cutting unsuitable for cutting lamination bevels. The key to plasma-are's successful usage in cutting bevels will be to solve the edge burning problem, minimize or even eliminate (if possible) the kerf taper, and to reduce the amount of dross.

The key to plasma-are's successful usage in cutting bevels will be to solve the edge burning problem, minimize or even eliminate (if possible) the kerf taper, and to reduce the amount of dross. Such an effort will require more than the optimization of process parameters which is beyond the scope of this paper.3.4 Laser CuttingLaser cutting is another promising non-contact method for machining bevels because of the rapid cutting speeds and steep bevel angles that can be achieved. The main issue to be resolved is what combination of laser type, beam temporal mode, and beam delivery system is best for cutting lamination bevels, particularly in steel. Therefore, a series of experiments were performed with various industrial laser types (C02 and Nd:YAG), cutting modes (pulsed and continuous wave), and beam delivery configurations (hard optics and fiber optics) to quantify the ability of a particular laser configuration to cut bevels into lamination material.

The object of these experiments was to determine what maximum bevel angles and cutting speeds were achievable.CO2 laser light can not be delivered by a fiber optic system. Since most glasses, including those used for fiber optic cables, have a high absorbtivity at the CO2 laser wavelength of 10.6 /lm, fiber optic delivery for a CO2 laser is not used and is not commercially available [Powell,1993]. A Light beam from a CO2 laser is almost always delivered with a hard optic system.CO2 laser cutting tests were performed using a Laser Dyne Model 780 5-axis CNC Machining System with a Rofin-Sinar 1700 SM laser. Using the laser fusion cutting method, a pulsed CO2 laser beam with hard optic delivery and nitrogen assist gas successfully cut bevels of up to 45° with narrow kerfs (0.18 to 0.30 mm) but with an extensive recast material (a.k.a. adherent dross) that is difficult to remove. Pulsing energies of2.7 to 6.5 joules were used. For reactive gas cutting (Le.

adherent dross) that is difficult to remove. Pulsing energies of2.7 to 6.5 joules were used. For reactive gas cutting (Le. oxygen assist gas) in pulsed and continuous wave (CW) modes, bevel angles of only 10° were achievable and most of the cuts exhibited self-burning, presumably from an excess of oxygen in the.1 'Self burning' is an uncontrolled oxidation reaction caused by an excess flow of oxygen.125the cut, and taper angle e. The control factors that affect these output responses are the focused spot diameter d which is proportional to the process lens focal length F, pulse width Tw, pulsing frequency /p, cutting speed Vc, type of assist gas (e.g. O2, Air, N2), and assist gas pressure Pg• All of these control factors (actual values used are listed in Table 3) are easily varied with the CNC laser cutter used for the experiments.

O2, Air, N2), and assist gas pressure Pg• All of these control factors (actual values used are listed in Table 3) are easily varied with the CNC laser cutter used for the experiments. Since almost all Nd:YAG lasers used for cutting have a tuned resonator, the output power of the laser beam is held constant to minimize the beam divergence. Therefore, the average laser output power was consistently held close to 0.25 kW in this case.Table 3 - Control factor values for laser cutting experiments.ex. Level %02 in assist gas (mass) PI> (kPa) F (cm) fn (Hz) V c (m/min) 'tw (msec)low 23140 7.6 30 0.13 0.30° med. -280 12.7 60 0.25 0.4high 100410 20.3 100 0.38 0.5low 23140 7.6 30 0.13 0.330° med. -210 12.7 60 0.25 0.4high 100280 20.3 100 0.38 0.5low 23140 7.6 22 0.13 0.360° med.

The kerf width w decreases with lower concentrations of oxygen, higher pulsing frequencies, shorter pulse widths, and shorter processing lens focal lengths. This last effect may be explained by the smaller effective focused beam diameter. No significant effects on kerf width were observed for changes in assist gas pressure or cutting speed. Less dross is noticed with higher oxygen concentrations and cutting speeds. The effects of assist gas pressure, processing lens focal length, laser pulsing frequency, and pulse width on HD are either negligible or unclear.In the following discussion, the parameter values corresponding to the low, medium, and high levels are found in :Table 3. For the smoothest surface finish, a low %02, medium fp, and high 'tw are recommended. For the narrowest kerf, a low %02, high fp, and low F are recommended. For the least amount of dross, a high %02, medium Pg , and high Vc are recommended.

For the narrowest kerf, a low %02, high fp, and low F are recommended. For the least amount of dross, a high %02, medium Pg , and high Vc are recommended. Depending upon what surface qualities of the bevel cut are most important, the levels of each laser cutting parameter can be set accordingly. The test bevel cuts with the best overall kerf quality and the associated process parameters are listed in Table 4.IfiTable 4 - OptImal contro factor eve s or anNd:YA aser cutter.GIex. 0° 30° 60°%02 High High HighPg High High HighControl Factor LevelsF Low Low Highfp High High Med.V c Med. Med. Low'tw High High Med.Output Responses Ra (!lm) w(mm) HD(mm) 0.14 0.10 0.070.02 0.04 0.392.6 4.4 5.9With optimal process parameters, lamination cutting with the same general surface finish as flute-edge endmilling and AWJ are possible but surface finish deteriorates with an increase in the bevel angle.

The edge burring due to adherent dross is comparable to AWJ but much less than endmilling. The kerf width is the smallest of all three bevel cutting methods. Although bevels up to ±80° are possible, the range of parameters that yields acceptable cuts narrows with the increase in bevel angle. Since there is a very small mechanical force associated with laser cutting, the fixturing required for lamination being cut is minimal. Furthermore, the HAZ of the laser cut is very small and most of the heat-affected material is blown away during the cutting process.126cut zone. However, the kerfs are narrow; the adherent dross is porous, brittle, and easily removable; and lower pulsing energies of only 0.7 joules were needed. Higher cutting speeds were achievable in the CW cutting mode.Initial Nd:YAG laser cutting tests were performed using a Lumonics JK701H laser with a fiber-optic delivery.

Higher cutting speeds were achievable in the CW cutting mode.Initial Nd:YAG laser cutting tests were performed using a Lumonics JK701H laser with a fiber-optic delivery. Using O2 as an assist gas, an Nd:YAG pulsed laser beam delivered with fiber-optics could cut bevels only up to 30°. Kerf widths are also narrow (e.g. 0.33 to 0.40 rom) but the cutting dross increases with the bevel angle. Cutting speeds were slightly lower than for a CO2 laser of similar power. Pulsing energies of 2.2 joules were used. Cutting in CW mode yielded unacceptable cuts.With the limited success of fiber-optic delivery, a Lumonics JK704 Nd:YAG laser with hard-optic delivery was used for cutting tests. When air (20% oxygen) was used as the assist gas for a pulsed Nd:YAG laser beam, bevel cuts up to 75° were achieved. Even higher angles may be achievable. Narrow kerf cuts (Le. 0.15 to 0.30 rom) were observed with a porous dross that remains relatively constant for all bevel angles.

Even higher angles may be achievable. Narrow kerf cuts (Le. 0.15 to 0.30 rom) were observed with a porous dross that remains relatively constant for all bevel angles. The lower cutting speeds are attributable to the low pulsing frequency used. Faster cutting speeds will be achievable at the low bevel angles (Le. 0 to 40°) if higher pulsing frequencies are used. For cutting experiments involving different lamination thicknesses and laser powers, the maximum cutting speed decreases non-linearly with increasing bevel angle because both the effective material thickness and laser energy absorbtivity ofthe material surface decreases.Summary: As previously mentioned, a high laser beam energy intensity at the material surface is of prime importance for cutting steep angle bevels. The limited success with the CO2 laser beam delivered with hard optics is attributable to insufficient energy intensity at these steep angles.

The limited success with the CO2 laser beam delivered with hard optics is attributable to insufficient energy intensity at these steep angles. The energy absorbtivity of steel and other metals to CO2 laser light is relatively low compared to Nd:YAG laser light, which contributes to this problem. The theoretical focused beam diameter is 0.102 rom in this case. The better success with Nd:YAG laser beam delivered with hard optics is attributable to the higher absorbtivity of this shorter wavelength light by the metal and a much smaller focused spot diameter, specifically 0.008 rom and 0.013 mm for the low and high powered lasers, respectively. A small focused spot diameter yields a high energy intensity.The limited beveling success of an Nd:YAG laser beam with fiber optic delivery is due to a large focused diameter and the homogenization of the laser beam within the fiber. The spot diameter of the laser beam delivered through a 1.00 rom diameter fiber is 0.5 rom.

The spot diameter of the laser beam delivered through a 1.00 rom diameter fiber is 0.5 rom. Even though the theoretical focused spot diameter is typically 2 to 5 times larger in reality [Powell,1993], the focused spot diameter from the hard optic delivery of an Nd:YAG laser beam is at least an order of magnitude less than the 0.5 rom spot. Furthermore, the laser beam exiting a fiber optic is homogenized as compared to the incoming beam. Homogenization lessens the exiting laser light's quality to a non gaussian spatial distribution leading to a lower energy density on the material surface.3.4.1 Optimization of Parameters for Pulsed Nd:YAG Laser Cutting with Hard-Optic DeliveryFor cutting bevels into steel sheet-the predominant material used for tooling-the optimal configuration is a pulsed Nd:YAG laser with hard-optic beam delivery. A pulsed laser beam creates a scalloped-type kerf (see Figure 6(a» by overlapping drilled holes (see Figure 6(b».

A pulsed laser beam creates a scalloped-type kerf (see Figure 6(a» by overlapping drilled holes (see Figure 6(b». The process parameters chosen for the previous laser cutting experiments yielded kerfs of generally poor surface finish and geometrical accuracy. In this section, the optimal parameters for better kerf quality from pulsed Nd:YAG laser cutting are determined.surface debris; 'v~ (al ~(b)exit conee= taper1\ ".II I,....

In this section, the optimal parameters for better kerf quality from pulsed Nd:YAG laser cutting are determined.surface debris; 'v~ (al ~(b)exit conee= taper1\ ".II I,.... d = mean hole diameter II IFigure 6 - (a) Top view ofa pulsed laser-cut kerf and (b) features oflaser-drilled holes.As shown in Figures 3 and 6, the easily measurable output responses that help to define the kerf quality of the laser cut bevel are the kerf width w, surface roughness of the kerf Ra, the height of the burr HD on the farside of1273.5 Comparison of Bevel Cutting MethodsBased on the previous discussion, some final comparisons can be made between flute-edge endmilling,abrasive water jet cutting and laser cutting (plasma-arc cutting not considered) for beveling steel laminations. •The higher cutting forces associated with flute-edge endmilling significantly deflect the unsupported portion of a lamination being beveled. AWJ cutting and laser cutting, i.e.

•The higher cutting forces associated with flute-edge endmilling significantly deflect the unsupported portion of a lamination being beveled. AWJ cutting and laser cutting, i.e. non-contact cutting methods, cause negligible deflection to the lamination. All three cutting methods are capable of high cutting speeds but the maximum speed is highly dependent upon the steel composition and hardness with flute-edge endmilling and less so with AWJ cutting. The maximum laser cutting speed is only dependent on laser power which means that higher cutting rates are achievable. The maximum bevel angle for all three methods is around ±80°. The kerf from AWJ cutting has a very large, consistent taper of around 10 0 for all bevel angles that must be compensated for during bevel cutting. Laser cutting creates only a slight kerf taper. Flute-edge endmilling leaves no appreciable kerf taper unless there is significant deflection of the lamination and cutting tool during beveling.

Laser cutting creates only a slight kerf taper. Flute-edge endmilling leaves no appreciable kerf taper unless there is significant deflection of the lamination and cutting tool during beveling. The width of the kerf affects the smallest radius of curvature achievable fora PEL's profiled edge. Of the three beveling methods, laser cutting creates the narrowest kerf. The kerf from AWl cutting is also narrow but not as much as that cut with a laser. The kerf width from endmilling is relatively large since it is the diameter of the cutting too1. Flute-edge endmilling leaves a very large machining burr, especially at larger bevel angles. AWJ cutting and laser cutting, on the other hand, yield much smaller machining burrs. Laser cutting with a pure oxygen assist gas yields a porous, brittle edge burr (i.e. iron oxide) which is easily removed from the cut lamination without grinding.

Laser cutting with a pure oxygen assist gas yields a porous, brittle edge burr (i.e. iron oxide) which is easily removed from the cut lamination without grinding. Flute-edge endmilling yields the smoothest surface finish of all the methods although the finish deteriorates at higher bevel angles from more machining chatter. AWJ cutting offers the most consistent surface finish for all bevel angles.•The author ranks the suitability (best to worst) of these three bevel cutting methods for cutting steel laminations as follows: I. Nd:YAG laser cutting because cutting speed is only dependent on laser power, tool wear is non-existent, cuttingforce is negligible and the kerf is the narrowest of all.2. Abrasive water jet cutting since cutting speed is dependent on material hardness and the kerf taper is the largest among the three methods.

Abrasive water jet cutting since cutting speed is dependent on material hardness and the kerf taper is the largest among the three methods. Flute-edge endmilling since the cutting force, kerf width, edge burr, and cutting speed's dependency on material hardness is the greatest overall.2. Abrasive water jet cutting since cutting speed is dependent on material hardness and the kerf taper is the largest among the three methods. Flute-edge endmilling since the cutting force, kerf width, edge burr, and cutting speed's dependency on material hardness is the greatest overall.4. Suitable Cutting Trajectory and Method for Cutting Steel PELsFrom the preceding discussion, uniformity in cut quality and geometry of steel laminations is based on the control of cutting process parameters. The main parameter that is controlled by the chosen cutting trajectory is cutting speed. In flute-edge endmilling, the maximum Vc is constrained by the steel's machinability and the lamination stiffness.

In flute-edge endmilling, the maximum Vc is constrained by the steel's machinability and the lamination stiffness. Furthermore, Ra deteriorates with increasing speed. For AWJ cutting, kerf width decreases with increasing speed but the effect of the water jet lag line becomes more pronounced. For laser cutting, surface finish gets better with increasing cutting speed but more extensive deburring is required. The ideal situation for cutting with any of these three of these cutting methods would be to cut lamination material at a constant speed for the entire cutting trajectory.While the methods of [Newman et al.,1995], [Zheng et al.,1996], and [Hope et al.,1997] can be used to define a trajectory for cutting the top edge of a PEL at a nearly constant velocity, the author has chosen a more straight-forward procedure.

As shown in Figure 7, the lamination's profiled edge is defined by two adjacent profiles resulting from the intersection of a two parallel cutting planes (separated by the lamination thickness) with the tool's CAD surface model (e.g., NURBS surface). Typically, each profile will be B-Spline curve, i.e., a piecewise polynomial made up of Bezier curves. To achieve a cutting trajectory with a nearly constant velocity, both curves are divided into the same number (n) of equal length segments using numerical procedures. Points are assigned to the end of each curve segment.

This series of corresponding point pairs (Pj,P2) are used to define corresponding line-of-site machining instructions, that is, a series of position point and unit directional vector pairs (P j, V) as If a constant velocity is prescribed for the point-to-point moves performed by the CNC shown in Figure 1(c).128machine (e.g., 5-Axis laser cutter), then the cutting means will have a nearly constant velocities along both profile curves.Farside B-Spline CurveP 2PIn5Figure 7 - Lamination's profiled-edge defined by two adjacent B-Spline curves.Acknowledgment This research was partially supported by the National Science Foundation CAREER award DMI-9734231.BibliographyBeyer, E., Klein, R., and Herziger, G., "Method for the Production ofThtee Dimensional Models," German Patent #P 37 II 470.0, Issued October 27, 1988.Fuchs, H., Kedem, Z.M., and Uselton, S.P., "Optimal Surface Reconstruction from Planar Contours," Communications ofthe ACM, Vol. 20, No. 10, Oct. 1977, pp.

20, No. 10, Oct. 1977, pp. 693-702.Glozer, G.R., and Brevick, 1R., "Laminate tooling for injection moulding," Journal ofEngineering Manufacture, Vol. 207, No. Bl, 1993,pp. 9-14.Hope, R.L., Jacobs, P.A., and Roth, R.N., "Rapid Prototyping with Sloping Surfaces," Rapid PrototypingJottrnal, Vol. 3, No.1, 1997, pp. 12 19.Keppel, E., "Approximating Complex Surfaces by Triangulation of Contour Lines," IBM Journal ofResearch Developments, Vol. 19, January 1975, pp. 2-11.Lee, C.H., Gaffney, T.M., and Thomas, c.L., "Soft Tooling for Low Production Manufacturing of Large Structures," Proceedings ofthe 1996 Solid Freeform Fabrication Conference, Austin, TX, August, 12-14, 1996, pp. 207-214.Matsui, S., Matsumura, H., lkemoto, Y, and Shimizu, H., "High Precision Cutting Method for Metallic Materials by Abrasive Watet:jet," Proceedings ofthe 6th American Water Jet Conference, Houston, Texas, Aug. 24-27, 1991, pp.

120, No.3, August, 1998.Weaver, W.R., "Process for the Manufacture of Laminated Tooling," US Patent #5031483, Issued ort July 16,1991.Zheng, Y, Choi, S., Mathewson, 8., and Newman, W., "Progress in Computer-Aided Manufacturing of Laminated Engineering Materia.ls Utilizing Thick, Tangent-Cut Layers," Proceedings ojthe 1996 Solid FreejormFabrication Conference, Austin, TX, August, 12-14, 355-362.129pp..

yond layer-by-layer Additive Manufacturing – Voxel-wise Directed Energy Deposition A. R. Nassar* and E. W. Reutzel* * Applied Research Laboratory at the Pennsylvania State University, University Park, PA 16802AbstractConventional additive manufacturing is a layer-by-layer process, reliant on the sequential deposition of 2-1/2 D layers oriented along a build axis. During directed energy deposition a feed- stock is directed into a continuous melt pool formed by a laser or electron beam. The ability to produce overhangs is limited due to the gravitational, surface tensions, and fluid-flow force acting on unsupported melt pools. Here, we present a novel, directed-energy-deposition technique where vertical and overhanging structures are formed by laser power modulation and the motion of a laser beam in three dimensional space along the build-up direction, rather than strictly in a single layer. We demonstrate that highly-overhanging Ti-6Al-4V structure, i.e.

We demonstrate that highly-overhanging Ti-6Al-4V structure, i.e. in which the overhang angle exceeds 45 degrees with respect to the x-y plane, can be deposited using the developed technique. High-speed imaging is used to gain insight into the physics of the process. The use of a pulsed or power-modulated beam is found to be critical to the formation of overhangs.1. IntroductionClassically-defined, additive manufacturing (AM) encompasses a family of technologiesbased upon successive deposition of 2-1/2 dimensional layers. Powder bed fusion (PBF) or directed energy deposition (DED) [1] can be used for fabrication of fully-dense metal components. In the former, a layer of powder is spread across a build plate or previously-fused area followed by the exposure of a rapidly moving heat source (e.g. electron beam or laser beam). The beam melts and fuses the powder to a portion of the underlying substrate.

electron beam or laser beam). The beam melts and fuses the powder to a portion of the underlying substrate. Directed energy deposition relies upon a different mechanism for the introduction of the feed-stock material: powder can be blown using a coaxial or off-axis nozzle arrangement into a melt pool formed by a heat source (e.g. electron beam, laser beam, or torch) or wire can be directed into the melt pool using a wire feeder.Overhanging structures—structures which do not lay entirely atop underlying, fusedmaterial—can be readily deposited using PBF processes because the surrounding powder in the bed supports overlaying material. This capability enables build-up of complex three-dimensional components. In practice, the overhang angle must not exceed a maximum overhang angle, which is dependent upon the material system and processing settings [2].

In practice, the overhang angle must not exceed a maximum overhang angle, which is dependent upon the material system and processing settings [2]. For example, a maximum overhang angle of approximately 65 degrees is suggested for EOSINT M systems [3].Figure 1: Overhang angle definitionThe maximum overhang angle for DED is more limited than for PBF. There is a paucityof literature regarding the maximum overhanging angle possible with DED, though the deposition of overhang angles of approximate 30 degrees has been claimed [4] and a 26 degrees overhang has been reported [5] using an Optomec LENS system with a conventional 3-axis stage arrangement and deposition head. Part reorientation via tilting or rotation of the substrate or deposition head can be used to extend the overhang angle [6], [7]; however, few software tools exist for automated planning of builds utilizing four or more axes of motion.

The use of robot arms may be an alternative but it’s use will be limited due to low line-following accuracy [8]. Additionally, it has been reported that data-transfer requirements for control of robotic arms results in build interruptions [7].Here, a method for deposition of highly-overhanging structures is developed. Unlikeconventional additive manufacturing, the method does not rely on sequential deposition of 2-1/2 dimensional layers; rather, voxels are sequentially deposited along a vector in the desired overhang direction. Deposition of overhanging structures with angles between 0o and 60o (as defined in Figure 1) are demonstrated.2. Experimental SetupASTM grade 5 titanium (Ti-6Al-4V) was used for both the substrate and feedstock powder. The powder was spherical, Extra Low Interstitials (ELI)-grade with a (-100+325) mesh size (44 to 149 𝜇𝑚). A directed-energy deposition, laser-based, Optomec, Inc. LENS MR-7 system was used for deposition.

A directed-energy deposition, laser-based, Optomec, Inc. LENS MR-7 system was used for deposition. The system utilized a 500 watt, Ytterbium-doped fiber laser with a second-moment beam spot size, measured using a Primes, GmbH FocusMonitor system, of 1.24 mm at the working distance. As shown in Figure 2, the working distance corresponded to a distance of 9.27 mm between the substrate and four, radially-symmetrically powder-delivery nozzles. Centered among the powder nozzles was a coaxial, center-purge nozzle, though which 30 lpm of argon flowed.During deposition, the processing chamber held an argon atmosphere at a gauge pressurebetween 2 and 3 inches of water (498-996 Pa), with oxygen concentration held below 20 partsper million. Argon gas was also used to assist the flow of powder through powder-delivery nozzles.

Argon gas was also used to assist the flow of powder through powder-delivery nozzles. Powder flow was measured and maintained at 3 grams per minute.Figure 2: Laser deposition arrangementOverhanging structures were formed by the translation of the substrate along a vectorrelative to the laser processing head. During translation, the laser power was modulated to produce a square wave with a peak power of 450 W, a period of 100 ms, and a 50% duty cycle. Unless otherwise stated, all deposits were formed at a constant vector translation speed of 0.42 mm/s (1 in/min) along a vector in a plane containing the laser propagation (build-up) direction. Vector was oriented at angles between 0 and 60 degrees with respect to gravity.During deposition, high-speed video was recorded using a Vision Research Inc.V1610 Phantom High Speed camera. The camera captured 1000 frames per second with an exposure time of 100 𝜇𝑠.

The camera captured 1000 frames per second with an exposure time of 100 𝜇𝑠. The camera was positioned outside the chamber, and viewed the process through standard laser-safe glass. A 200 mm focal-length lens was used with an F-stop of 32.Deposited samples were sectioned, ground, and polished according to standardmetallographic techniques. Samples were etched using Kroll’s reagent prior to imaging.3. Results and Data AnalysisOverhanging structures were successfully deposited between 30 and 90 degrees with respect the horizontal (x) axis. The deposited structures are shown in Figure 3. Overhanging deposit lengths as long as 50.8 mm (2 in) were deposited.

The deposited structures are shown in Figure 3. Overhanging deposit lengths as long as 50.8 mm (2 in) were deposited. A 50.8 mm long deposit with an overhang angle of 45 degrees is shown in Figure 4.Figure 3: Overhanging structures deposited via voxel-wise depositionSeveral characteristic features of the overhanging deposits are visible in Figure 3 andFigure 4: Near the substrate, the diameter of the deposit is approximately half of the steady-state deposit width. The steady-state deposit diameters, approximately 5 mm in all cases, was more than four times greater than the beam spot size. The downward-facing surface appeared smoother than the upward-facing surfaceMetallographic examination revealed the formation of columnar prior beta grains oriented along the overhang build-up direction, rather than vertically with respect to gravity as typically found in layer-by-layer builds.

As in other AM techniques, the prior-beta grains appear to have grown epitaxially starting from previously deposited material [9]. Within the prior-beta grains, a fine basketweave alpha-beta microstructure was observed. This is similar to other AM- deposited Ti-6Al-4V microstructures.Figure 4: Forty-five degree overhanging structure.Figure 5: Micrographs and macrograph of a Forty-five degree overhanging structure.High-speed imaging provided insight into the process by which the overhanging structures were formed. High-speed images captured at the beginning of the process during the initial deposition onto the substrate are shown in Figure 6. With respect to the axes shown in the figure, the structure was built up along the y-z plane toward the camera. At the beginning of the deposition, during the first pulse period, a semi-spherical melt pool formed. Within 10 ms, a hemispherical deposit was observed extending from the substrate.

At the beginning of the deposition, during the first pulse period, a semi-spherical melt pool formed. Within 10 ms, a hemispherical deposit was observed extending from the substrate. The build continued to grow throughout the laser-on cycle. Accompanying the deposition process was a metal-vapor plume. The plume maintained a parabolic shape, but fluctuated in length, width, and angle of emergence.The laser-off cycle resulted in rapid cooling of the melt. Based on close-up examination of flow within the melt pool, it was estimated that full solidification of the first voxel occurred within 25 ms of the laser beam turning off.Figure 6: High speed images of the first lase-pulse period during deposition of a 45 degree overhanging structure. At the bottom of each frame, the frame time is displayed.Solidification during the laser-off cycle is critical to the formation of the highly-overhanging structures.

At the bottom of each frame, the frame time is displayed.Solidification during the laser-off cycle is critical to the formation of the highly-overhanging structures. The results of an attempt to build a 45 degree overhanging structure using a continuous (CW) laser beam power of 450 W, with all other variables equal, shown in Figure 8, illustrates the importance of a solidification cycle.Figure 7: High speed images of a laser pulse part-way (26.3 seconds) through the build. At the bottom of each frame, the frame time, with respect to the laser-on trigger, is displayed.Using a CW beam, an overhanging structure did not form. High-speed imaging of theCW deposition attempt over the course of 30 seconds (Figure 9) revealed the formation and continued growth of a large, melt pool held together by surface tension. The melt pool increases in size throughout the deposition process. As the deposition head moved within the y-z plane, the rear of the melt pool solidified.

The melt pool increases in size throughout the deposition process. As the deposition head moved within the y-z plane, the rear of the melt pool solidified. The melt pool solidified approximately 174 ms after the laser beam was turned off—this is determined based on observation of flow within the melt pool— leaving behind a globule laying atop the substrate.Figure 8: Attempted build-up of 45 degree overhanging structure using a CW laser beam.Figure 9: High-speed imaging of attempted build-up of 45 degree overhanging structure using a CW laser beam.4. DiscussionLaser power modulation is critical to the formation of highly-overlapping structures bytranslation along an overhang vector. When the laser is on, a melt pool begins forming below the laser beam. This molten zone, held together by surface tension and acted upon by gravity, expands spherically while the laser beam is on.

When the laser is on, a melt pool begins forming below the laser beam. This molten zone, held together by surface tension and acted upon by gravity, expands spherically while the laser beam is on. When the laser beam extinguishes, the melt rapidly solidifies—starting at the bottom and sides of the pool. The process repeats as the focal position of the powder and the laser spot translate along an overhang vector. Thus, the formation of overhanging structures is aided by two physical phenomena: surface tension coaxes the pool to form a spherical shape; and, the solidification of the melt pool exterior acts like a vessel for the solidifying voxel.The formation of highly-overhanging structures is not possible with a CW laser beam.Continued heating of the melt pool results in the formation of large spherical melts, held together by surface tension. Without solidification of the sides of the melt, the upward build-up becomes impossible.

Without solidification of the sides of the melt, the upward build-up becomes impossible. Critical to the process is tuning of the upward motion of the stage and the laser pulse parameters: upward motion at a rate exceeding the spherical growth of the melt will lead to failure. Upward motion at a rate less than the special growth of the melt will result in a loss of process efficiency and poor geometric conformity. Within this limit, tuning of the processing parameters allows for variation in the diameter of the overhanging structures5. Concluding RemarksThe presented method for the formation of highly-overhanging structures is a significantadvancement for directed-energy deposition processes. Key to the process is the pulsing of the laser beam to allow solidification of partially-underlying voxels. However, complete solidification of an underlying voxel is not a requirement; solidification of the voxel walls are sufficient to prevent collapse of the melt pool.

However, complete solidification of an underlying voxel is not a requirement; solidification of the voxel walls are sufficient to prevent collapse of the melt pool. Given this, alternative arrangements, where beam spot size, intensity distribution, power, or powder flowrate is modulated, may be capable of producing similar results.The process may be extended to build-up of fully-dense, three-dimensional structures,such as the simple structure illustrated in Figure 10, as well as the formation of supports. Further development of the process, together with numerical modeling, and the development of process and path planning software, may result in the ability to deposit complex, three-dimensional structures that are currently only feasible with powder-bed fusion technology, five-axis motion system, and robotic arms.

Another key application of the voxel-wise deposition may be in the deposition of large, cellular structures, such as that illustrated in Figure 10.Figure 10: (left) Illustration of motion paths for the deposition (middle) of a simple three- dimensional structure using voxel-wise deposition. (right) Illustration of a cellular structure which may be deposited using voxel-wise deposition.AcknowledgementsThe authors gratefully acknowledge the assistance of Joseph Welz in performing high-speed imaging of the laser deposition process. This work was supported by the Office of Naval Research, under Contract No. N00014-11-1- 0668. Any opinions, findings and conclusions or recommendations expressed in this publication are those of the authors and do not necessarily reflect the views of the Office of Naval Research.References[1] ASTM F42 Committee, “Terminology for Additive Manufacturing Technologies,” ASTMInternational, 2012.[2] D. Wang, Y. Yang, Z. Yi, and X.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperBIG AREA ADDITIVE MANUFACTURING APPLICATION IN WIND TURBINE MOLDSBrian K. Post, Bradley Richardson, Randall Lind, Lonnie J. Love, Peter Lloyd, Vlastimil Kunc, Breanna J. Rhyne, Alex Roschli, *Jim Hannan, *Steve Nolet, *Kevin Veloso, **Parthiv Kurup, ** Timothy Remo, and **Dale JenneManufacturing Demonstration Facility, Oak Ridge National Laboratory, Knoxville, TN 37932, USA *TPI Composites, Inc. PO Box 367, 373 Market Street, Warren, RI 02885, USA ** National Wind Technology Center, National Renewable Energy Laboratory, Boulder, CO 80303, USAAbstractTooling is a primary target for current additive manufacturing (AM), or 3D printing, technology because of its rapid prototyping capabilities. Molds of many sizes and shapes have been produced for a variety of industries.

Molds of many sizes and shapes have been produced for a variety of industries. However, large tooling remained out of reach until the development of large-scale composite AM manufacturing processes like the Big Area Additive Manufacturing (BAAM) system. The Department of Energy’s Oak Ridge National Laboratory (ORNL) worked with TPI Composites to use the BAAM system to fabricate a wind turbine blade mold. The fabricated wind turbine blade mold was produced in 16 additively manufactured sections, was 13 meters long, had heating channels integrated into the design, and was mounted into a steel frame post fabrication.

This research effort serves as a case study to examine the technological impacts of AM on wind turbine blade tooling and evaluate the efficacy of this approach in utility scale wind turbine manufacturing.IntroductionThe purpose of this paper is to identify potential use of Additive Manufacturing (AM) in theproduction of composite parts for wind energy systems.AM, more commonly referred to as 3D printing, is a fast-growing industry in which recenttechnological advances have expanded its use beyond rapid prototyping. AM encompasses techniques applicable across multiple material types—most commonly polymers, metals, composites, and ceramics—that use computer-rendered designs to produce a near-net-shape part, layer by layer.

Current costs for both polymer-based and metal-based AM systems keep AM out of large-volume production, but advancements in AM techniques will lower per-unit costs and advance designs beyond the capabilities of current conventionally manufactured (CM) components.In the wind industry, current and research and development-level (R&D) AMtechnologies have potential for considerable impact on the prototyping and manufacturing cost of wind energy tooling and components. In the future, AM technologies could enable on-site manufacturing of turbine components as well as production of site-optimized components that are tailored to the unique wind and grid resources of a given location. For example, lower-cost2430and faster production of blade molds can increase the rate at which industry prototypes and conducts full-scale field testing of novel designs. Easier prototyping opportunities could be extended to customized nacelle designs and turbine configurations.

Easier prototyping opportunities could be extended to customized nacelle designs and turbine configurations. AM could be used to create specialized tooling for the manufacturing of conventional components. AM will also allow for the development of components designed to better suit the specifications of their use, including lower materials requirements.Insights gained from this project can be applied to the manufacture of other components forwind energy systems. Additional analysis of AM’s current and newly-designed wind system components, the use of alternative materials, and on-site manufacturing potential is recommended. Rapid development of polymer and metal AM capabilities continue to cause costs to decline, which makes finding break-even points of AM and CM parts a constant moving target.

Rapid development of polymer and metal AM capabilities continue to cause costs to decline, which makes finding break-even points of AM and CM parts a constant moving target. Future R&D and analysis will further identify cost reduction opportunities and limits to AM applications.Big Area Additive ManufacturingConventional fused deposition modeling is based on melting and extruding a filament of thermoplastic feedstock. Prior work shows that the peak flow rate is limited by the rate at which the filament can be melted [1]. For larger print volumes (at least 15 times the build volume of desktop-size consumer-level AM systems), technologies are developing. For example, Cincinnati Incorporated partnered with ORNL to developed the Big Area Additive Manufacturing (BAAM) process in 2014, which scales the extrusion process from desktop-sized AM systems of 1-5 cubic inches per hour to over 1000 cubic inches per hour [2].

To achieve high print rates, material pellets are used, which also significantly lowers the material cost (from $110 – $220/lb to $3 – $5 / lb), in combination with an optimized feed screw that reduces throttling of the pellet flow into the material. BAAM is an extrusion process that uses injection molding material for the feedstock and a single screw extruder for melting and metering the flow rate (see Figure 1) [3]. A gantry system, commercialized by Cincinnati Incorporated (see Figure 2), moves the extruder in x, y, and z directions to build the part. The extruder is capable of delivering 100 lbs/hour of thermoplastic materials from pellet feedstock. The gantry system is capable of achieving 200 inch/sec peak velocities with 64.4 in/s2 accelerations and position accuracy in the 0.002”. The use of carbon fiber reinforcement in the thermoplastic resins increases the part strength and stiffness [4].

The use of carbon fiber reinforcement in the thermoplastic resins increases the part strength and stiffness [4]. Just as important, carbon fiber reinforcement also increases thermal conductivity and reduces the coefficient of thermal expansion lessening the need for a heated chamber to produce large parts (see Figure 3 and Figure 4) [5]. The elimination of the oven significantly decreases the energy intensity, which is the manufacturing energy required per kilogram of product. As shown in Figure 5, conventional FDM systems with heated chambers have a 100 kW-hr/kg energy intensity. Desktop systems that have similar production rates have a 5.5 kW- hr/kg energy intensity suggesting that the oven accounts for 95% of the energy consumption in FDM production. BAAM further reduces the energy intensity to 1.1 kW-hr/kg by significantly increasing productivity from 1 ci/hr to 2500 ci/hr, and BAAM parts are manufactured at room temperature.

BAAM further reduces the energy intensity to 1.1 kW-hr/kg by significantly increasing productivity from 1 ci/hr to 2500 ci/hr, and BAAM parts are manufactured at room temperature. Due to its higher build rates and build sizes, BAAM’s initial targeted applications is in the tooling industry.

Due to its higher build rates and build sizes, BAAM’s initial targeted applications is in the tooling industry. Specifically, in large, complex injection molds that are needed to produce CM polymeric and metal parts at a low cost.2431Figure 1: BAAM ExtruderFigure 3: Section of wind turbine mold2432Figure 2: Cincinnati BAAMFigure 4: Printed prototype houseFigure 5: Energy intensity for manufacturingTherefore, the combination of lower energy intensity, higher productivity, and lower material cost suggests that there will be a significant production cost reduction with BAAM when compared to conventional AM systems.Applications of Current Additive Manufacturing Systems in Wind Power Systems: Fabricating Blade MoldsAM is hindered by limited production rates, expensive materials, processing costs, and small build volumes. These conditions serve to significantly limit the applicability of AM to wind energy system production.

These conditions serve to significantly limit the applicability of AM to wind energy system production. Commercial turbines are large systems where individual components often exceed 50 meters in length. The sheer size and mass of these components combined with throughput requirements largely preclude the application of current AM technology in wind power manufacturing.However, several application areas have been identified where economic and technological incentives make AM processes viable for current production. Furthermore, if wind applications were used as a target to drive the development of AM systems, more applications with significant economic benefits become achievable.The first identified wind application for AM, and the target of this study, is in the manufacture of molds for wind turbine blades.

As described previously, tooling is a primary target for the current state of AM technology given the less stringent requirements for material2433performance and certification. AM molds of various sizes and shapes have been produced for a variety of industries. Large tooling, however, remained out of reach until the development of large scale composite AM processes like the BAAM-CI. Figure 6 shows the ORNL/TPI AM blade mold demonstrator produced as a collaboration between ORNL and TPI and funded by DOE’s Advanced Manufacturing and Wind and Waterpower Offices.Figure 6: Additively manufactured blade mold and produced blade section.The traditional blade mold manufacturing process is a multi-step, expensive operation. First, a plug is manufactured by subtractive machining of foam and tooling resin. This multi- piece assembly takes several weeks to months to produce from CAD drawings that are then used to generate CNC toolpaths.

This multi- piece assembly takes several weeks to months to produce from CAD drawings that are then used to generate CNC toolpaths. From there, the plug is shipped to the mold manufacturer where it is qualified and aligned. The mold manufacturer then applies a release agent and lays up fiberglass of sufficient thickness to support the molding operations. Miles of heating wire is manually embedded in the underside of the mold and arranged in zones of similar area. Then a steel frame is manually erected and attached to the mold prior to removal from the plug. From one plug set, up to eight molds can be produced amortizing the cost. For the production of one prototype mold, or a small set of molds, this cost can be prohibitively expensive, often reaching millions of dollars.ORNL- TPI Scaled Wind Farm Technology (SWiFT) Facility MoldTo study the effect of wakes in large wind farms, the Sandia SWiFT facility uses a scaled wind farm.

These turbines require custom blades to simulate the dynamics of their larger counterparts. The blades are scheduled to be produced using additively manufactured molds produced in partnership between ORNL and TPI and funded by the Department of Energy’s (DOE) Advanced Manufacturing (AMO) and Wind and Water Power Office (WWPO). This research effort serves as a case study to examine not only the technological impacts of AM on2434wind blade production, but also the economics of AM at the large scale using BAAM technology.Mold Requirements and Risk Reduction ExerciseThe first step to the project was to define the mold requirements. For the resin infusion, there were a number of critical performance targets for the mold necessary to successfully manufacture blades. These targets included operating temperature, temperature gradients across the surface of the molds during service, mold distortion, vacuum, and life. The team identified three goals for each metric.

These targets included operating temperature, temperature gradients across the surface of the molds during service, mold distortion, vacuum, and life. The team identified three goals for each metric. The first was the target necessary for the success of the project. The second was a stretch goal that, if passed, would qualify the process for low volume production. Finally, the ultimate goal for the technology was to meet the targets necessary for full scale production. The metrics and goals for the project are outlined in Table 2.Table 2: Project Metrics and GoalsThe first phase of the project focused on risk reductions steps. The team designed a series of experiments to ensure that the BAAM molds would achieve the targets. High risk items included the heating (achieving the target temperatures and temperature gradients), distortion, and vacuum integrity. In terms of vacuum integrity, the team elected to coat the printed mold with fiberglass.

In terms of vacuum integrity, the team elected to coat the printed mold with fiberglass. The mold was designed 4 mm under the target mold lines. TPI coated the molds with 8 mm of fiberglass that was then machined to the final surface geometry and finish. The second item was the heating of the mold. It was determined that additive manufacturing could enable a very novel approach to heating. Rather than manually inserting miles of conductive heating wires, the molds could be designed with ducting directly printed into the structure. Small heating elements and blowers could be easily and rapidly installed in the molds, taking the time for installation of heaters from weeks and even months to a single day.2435Table 2: Performance Metrics and GoalsFigure 7: Integrated ORNL/TPI heater units in blade mold sectionFigure demonstrates the arrangement of heater units in the SWiFT mold set for the uniform heating of the mold surface.

Each channel has been designed to be of equal surface area to reduce non-uniformity of the temperature profile over the surface of the mold. Hot air is circulated through the mold until the surface temperature reaches the desired set point; then the set point is lowered to a maintenance level.Figure 8: Airflow in mold sectionA cross sectional view of the second mold section is shown in Figure 8. Arrows illustrates the heat flow, in that hot air exits the heater unit and flows along the surface of the mold where it loses energy and recirculates forming a closed loop. The functional structure of the AM component is used not only to achieve the desired surface geometry, but also to provide a housing for the heating unit and to distribute the heat uniformly throughout the structure. AM, in this example, enables functional structure and eliminates the manual labor associated with the emplacement of the heating system.

AM, in this example, enables functional structure and eliminates the manual labor associated with the emplacement of the heating system. Furthermore, a failed heating unit is easily replaceable via a pocket on the underside of the return channel.2436A single section was manufactured and evaluated for performance (see Figure 9). A series of experiments were conducted to evaluate the performance (distortion, temperature gradient, vacuum integrity) of the mold. A laser tracker (see Figure 10) profiled the accuracy of the mold at room temperature. The heaters were turned on and the system was then brought up to the target temperature. At target temperature, the laser scanner measured the hot surface profile, and thermal imaging cameras measured the temperature gradients across the surface (see Figure 11, Figure 12, and Figure 13). The temperature gradient was measured at +/- 3C at 40C exceeding the requirements for the stretch goal.

The temperature gradient was measured at +/- 3C at 40C exceeding the requirements for the stretch goal. As shown in Figure 14, surface variations are under +/-0.025”, approximately 0.1% of the chord. In terms of vacuum, the system exceeded the target of 15 mbar over 60 minutes. Therefore, the risk reduction exercise demonstrated that the AM mold sections exceeded the stretch goal requirements.Figure 9: Test sectionFigure 10: Laser tracker2437Figure 11: Thermal imaging cameraFigure 12: Output of thermal imaging camera and thermocouples2438Figure 13: Thermocouple outputFigure 14: Measured surface variation at molding temperaturePrinted Mold Manufacturing After successfully achieving the goals through the risk reduction exercise, the project transitioned to the development of a full scale 13 m blade mold. From a manufacturing perspective, the mold proved to be better suited for vertical printing (see Figure 15).

From a manufacturing perspective, the mold proved to be better suited for vertical printing (see Figure 15). The work volume in the BAAM system was 8 ft wide, 20 ft long, and 6 ft in height. Therefore, each mold was made from eight segments (see Figure 16) , printed one at a time except for the smaller tip sections which were printed in twos.2439Figure 15: Printed low pressure side section 1 in the BAAM CI printerFigure 16: Mold sections2440Figure 17: ORNL/TPI AM blade mold conceptual modelEach section is printed on end to maximize the resolution of the BAAM process and reduce the required support material as shown in Figure 15. Figure 18 shows the sections of the SWiFT AM blade mold printed prior to covering it with fiberglass.

Figure 18 shows the sections of the SWiFT AM blade mold printed prior to covering it with fiberglass. While many of the technical opportunities of AM in wind mold production are addressed in the production of the SWiFT mold, it also serves as a tool for the evaluation of economic incentives for similar molds on larger scales.Figure 18: 5 of the 8 low pressure side SWiFT mold sections2441After completion of printing, the molds were coated with fiberglass (see Figure 19) and machined to the final target surface (see Figure 20). The molds were fitted to be inserted into the egg crate structure (Figure 21) and were finally polished for operation (Figure 22).

The molds were fitted to be inserted into the egg crate structure (Figure 21) and were finally polished for operation (Figure 22). As of the writing of thispaper, one full set of blades has been manufactured (see Figure 23, three in total) with no discernable wear on the molds.Figure 19: Fiberglass surfaced printed componentFigure 20: Section being machined2442Figure 21: Assembly in egg crateFigure 22: Final molds (high pressure and low pressure)2443Figure 23: Finished bladeConclusions and Future ApplicationsAM is finding more and more industrial applications in the area of molds, jigs, and fixtures. The automotive, aerospace, and appliance industries are rapidly transitioning to printed tools because of the complexity, high cost, and long lead times associated with these parts.In terms of scale, there are new equipment manufacturers developing systems that are oneto two orders of magnitude larger and faster than the current abilities of the Cincinnati BAAM.

Ingersoll Machine Tool company announced the partnership with ORNL to develop the Wide High Additive Manufacturing (WHAM) system, see Figure 24 and Figure 25. The system will have a build volume that is approximately 25 feet wide, 20 feet tall, and 100 feet long and is scalable to much larger dimension. The production rate will start at 1000 lb/hr but is scalable to higher rates. This scale system enables the direct manufacture of parts (molds, blades, nacelles, etc.) that are competitive with current wind manufacturing needs.2444Figure 24: Front view WHAMFigure 25: Rear view WHAMGiven significant leaps in material properties and out of plane printing, there is potentialto directly manufacture wind turbine blades (see Figure 27). First, from the materials perspective, everything used to date in the BAAM process is based on short chopped carbon fiber.

First, from the materials perspective, everything used to date in the BAAM process is based on short chopped carbon fiber. These materials are ideal for tooling because of their high stiffness, but they do not have the strength needed for strength limited structural applications. There will be a need for long or continuous fiber reinforcement to directly manufacture wind turbine blades. In addition, to keep part weight down, it will be necessary to print core structures (such as foam) to achieve specific strength and stiffness requirements. Systems such as the WHAM will have the ability to print, machine, and coat all within one machining center.

Systems such as the WHAM will have the ability to print, machine, and coat all within one machining center. At rates of 1000 lb/hr and greater, it is feasible to envision rapid automated manufacturing of customized wind turbine blades with multi-material (carbon fiber, glass fiber, foam, etc.) solutions.Figure 26: 3D printed wind turbine blade2445AcknowledgementsThis work was sponsored in a joint partnership by the Department of Energy’s EnergyEfficiency and Renewable Energy’s Advanced Manufacturing Office and Wind and Water Power Technology Office. Cost share was provided by TPI composites in accordance with the Cooperative Research and Development Agreement. Matthew Sallas was integral to the fabrication and machining of the printed mold segments and this work would not have been possible without his hard work.References[1] Monzon, M., , Gibson, I., Benitez, A., Lorenzo, L., Hernandez, P., Marrero, M. (2013).

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperBIMETALLIC CASTINGS FOR WEAR PERFORMANCE THROUGH INFILTRATION OF ADDITIVE MANUFACTURED METAL LATTICE STRUCTURESJ. C. Liggett1, D. A. Snelling1, M. Xu2, O. J. Myers3, and S. M. Thompson41Department of Manufacturing Engineering, Georgia Southern University, Statesboro, GA 304602Department of Mechanical Engineering, Georgia Southern University, Statesboro, GA 304603Department of Mechanical Engineering, Clemson University, Clemson, SC 296314Alan Levin Department of Mechanical and Nuclear Engineering, Kansas State University, Manhattan, KS 66506AbstractHigh chromium white iron is an alloy frequently employed in the production of abrasion resistive wear components. Ground engaging components for mining or earthmoving frequently require such materials, as well as slurry pumps for mining applications.

Ground engaging components for mining or earthmoving frequently require such materials, as well as slurry pumps for mining applications. Although high chromium white iron alloy demonstrates excellent wear performance due to the formation of chromium carbides, it is brittle and lacks toughness. Impact resistance is often of great importance for ground engaging wear components; hence, this study will investigate a method by which high chromium white iron wear components may be reinforced by the formation of a bimetallic composite. In this research, an additively manufactured lattice structure of 316L stainless steel is infiltrated with high chromium white iron via the metal casting process. This procedure results in a bimetallic casting of reinforced white iron.

This procedure results in a bimetallic casting of reinforced white iron. Complete infiltration and metallurgical diffusion bonding were observed between the two alloys, validating this method as a means of reinforcing high chromium white iron castings for applications requiring high abrasion and impact resistance.IntroductionAlloy selection for metal components is often a compromise between desired material properties which are typically mutually exclusive. Applications requiring both high toughness and high wear resistance in metal alloy components often lead to significant compromise, as high hardness is advantageous for wear resistance but can result in a brittle material with low toughness [1].

Because desired material property combinations can be mutually exclusive within most engineering alloys, much development has been done recently to develop bimetal composites which can blend the strengths of two different alloys to make up for the weaknesses of each.The majority of the research and development in this area has been focused on bimetal casting techniques, in which conventional metal casting techniques are used to produce a composite through pouring successive layers of different molten alloys into a mold cavity, or by pouring a liquid metal alloy onto a solid metal substrate. [2–12] However, little has been done to study the applications of advanced additive manufacturing techniques in the development of bimetal castings.

[2–12] However, little has been done to study the applications of advanced additive manufacturing techniques in the development of bimetal castings. This study will seek to fill this gap by examining the feasibility of creating a bimetal composite through a hybrid casting-additive manufacturing process.742Existing work has been performed exclusively examining the casting process in the development of bimetallic wear components. In separate studies performed by Zic et al. and Zhu et al., the viability of bimetallic casting techniques involving the pouring of successive layers of liquid metal alloys is explored. [4, 5] In both cases, a wear resistant white iron alloy is coupled with a ductile steel to improve the impact response of the casting. In another study, Wróbel examines the feasibility of producing a bimetallic composite by pouring a liquid metal alloy over a solid metal substrate.

In another study, Wróbel examines the feasibility of producing a bimetallic composite by pouring a liquid metal alloy over a solid metal substrate. [6] These studies each provide important insight into the application of metal casting techniques toward bimetallic composite production.While much has been done to examine the production of bimetal composites through metal casting techniques, little work has been performed examining the applications of additive manufacturing in further developing bimetallic composite casting. Additive manufacturing can be used to create structures of high complexity with maximum material efficiency, low tooling cost, low labor cost, and low environmental impact [13].

Additive manufacturing can be used to create structures of high complexity with maximum material efficiency, low tooling cost, low labor cost, and low environmental impact [13]. Additive manufacturing (AM) of metal feed stock materials in particular has positively impacted a multitude of industries, being capable of producing complex end-use components with good surface finish without the complex processes involved in traditional manufacturing methods such as metal casting. AM has also been applied recently to the production of bimetallic composites in order to produce components of different compositions through a single manufacturing process.

AM has also been applied recently to the production of bimetallic composites in order to produce components of different compositions through a single manufacturing process. The thermal mismatch between materials in bimetallic AM has been overcome in some cases by a composition gradient in the transition zone in powder bed fusion processes, a process in which a metal feedstock powder is bonded through a laser or electron energy source to produce a final structure [13].While new AM methods have provided the potential for bimetal component production, they fall short in a few key areas. First, metal AM components are limited by deposition rates and therefore cannot scale up beyond a certain point [14]. Thermal management is tied intrinsically to deposition rates and thus there is a maximum print volume in which powder bed fusion methods can operate successfully [14].

Thermal management is tied intrinsically to deposition rates and thus there is a maximum print volume in which powder bed fusion methods can operate successfully [14]. While other metal AM methods such as directed energy deposition offer the potential to overcome this size limitation, materials manufactured through such methods possess inherent microstructure flaws and crack defects which do not suite the material to high toughness or wear applications [14,15].To best consider how AM methods may be used to improve bimetallic metal casting, it is important to examine areas in which AM techniques excel. One application for which AM is ideally suited is the production of complex lattice structures. Lattice or cellular structures can be seen throughout nature, featuring high strength, high stiffness, better energy absorption, and superior thermal and acoustic properties.

Lattice or cellular structures can be seen throughout nature, featuring high strength, high stiffness, better energy absorption, and superior thermal and acoustic properties. [16] While these structures are desirable in many engineering applications, traditional manufacturing processes struggle to produce such complex geometric structures or incur high costs during production [16,17]. AM however excels at the production of complex cellular structures, and powder bed fusion techniques in particular offer the ability to produce high strength metal lattice structures with minimal production challenges [17,18].Lattice structures excel at high strength to weight ratio applications because they are designed by placing material only where it is necessary [19]. Such structures can be designed as743either stochastic structures, which feature a randomly distribution of voids, or periodic structures, which feature ordered, repeated unit cell structures [17,19].

Periodic structures are often easier to design and control, as the overall shape can be broken down into predefined structural unit cells sometimes referred to as Manufacturable Elements [19]. These structural unit cells can be combined easily using element approximation methods to turn a solid part design into a lattice composed of unit cells of known structural properties, and thus the mechanical behavior of the lattice can be estimated [19]. Furthermore, lattice structures can be of uniform composition or can be hierarchical, in which the strut and cell dimensions change with location in the model. Examples of hierarchical lattice structures are often found in nature and are shown to improve overall structural performance [20].In a study performed by Zhong et al., the implementation of periodic unit cell lattice structures in the production of lightweight structural components was examined.

It was shown that the use of lattice structures was capable of reducing weight by 62%, compared to 54% reduction made possible by topology optimization alone [21]. Zhong et al. studied common cubic unit cell lattices, tetrakaidecahedron, diamond, and body-centered cubic structures. These shapes were composed of cylindrical struts, with study parameters varying with unit cell length, strut diameter, and lattice volume fraction. It was found that, given similar volume fractions and unit cell sizes, tetrakaidecahedron structures demonstrated significantly superior mechanical properties compared to diamond or body centered cubit structures. As found by this study, unit cell lattice structures offer key advantages over general topology optimization and special unit cell structures are ideally suited for high strength to weight ratio applications.More complex lattice structures can be modeled through relative density mapping methods [22].

According to Alzahrani et al., the overall strength of a structural lattice is dictated by the volumetric density of the lattice [22]. The strengths of a lattice can be evaluated by calculating the deformation of the interconnecting strut network in the lattice. For this reason, when the volumetric density of the lattice is higher, more or thicker struts will be present which thereby increases the strength of the structure [22]. In Relative Density Mapping (RDM), the strength of the lattice is calculated by the volumetric density of finite elements within the lattice using a mathematical formula to correlate the strut size with each element’s relative density.Further advanced design techniques proposed by Wang et al. address limitations in lattice design [23]. As noted in this study, most design techniques do not evaluate the global optimized solution for the lattice.

address limitations in lattice design [23]. As noted in this study, most design techniques do not evaluate the global optimized solution for the lattice. It is noted that determining the true global optimization is difficult due to disparity between the lattice structure distribution and the stress distribution in the structure and due to anisotropic material properties inherent in AM manufacturing techniques [23]. In this study, a method was developed which allowed for global optimization of a lattice structure through consideration of the principal stress lines through the component. Principal stress lines, orthogonal curve pairs which delineate the internal stress trajectory through a model, are shown to represent and support loading more efficiently than a grid-based approach [24].

A lattice structure is then created which conforms to the principal stress lines in the model.Applications of AM lattice structures are numerous; with most structures being built using the laser-powder bed fusion (L-PBF) AM method. Zhao et al. demonstrated the ability of Ti-6Al- 4V alloy lattice structures to absorb energy in high impact applications [25]. Gyroidal lattice744structures were shown to exhibit layer by layer failure in compression studies which allowed the load-bearing capability of the structure to increase during compression. Other studies show the ability of AM lattice structures to form lattice scaffolds for applications in tissue engineering [26] [27,28]. Large scale lattice structures for high strength to weight ratio structural applications were explored in a study performed by Yan et al. [18].

Large scale lattice structures for high strength to weight ratio structural applications were explored in a study performed by Yan et al. [18]. Selective laser melting AM methods were capable of producing gyroid lattice structures with unit cell size ranging up to 8mm.While lattice structures produced through AM show significant applications in engineering design, such structures are not without draw-backs. In the same study which showed the viability of lattice structures with unit cells ranging up to 8mm in size, the yield strength and elastic modulus was observed to decrease with increasing unit cell size, indicating an upper limit to lattice unit cell sizing [18]. Additionally, the presence of porosity defects and rough surfaces in components made by metal AM methods is unavoidable, and can lead to poor fatigue performance [29,30]. Delamination defects are another inherent weakness of AM methods which can critically reduce material properties of AM components [31,32].

Delamination defects are another inherent weakness of AM methods which can critically reduce material properties of AM components [31,32]. For this reason, there is significant room remaining for improving AM lattices for use in large scale structural applications.In this study, the development of a bimetal composite is proposed which incorporates the recent development of AM lattice structures with the metal casting process. The L-PBF AM method will be used to produce a tetrakaidekahedron lattice structure of 316L Stainless steel with superior strength to weight ratios [16–18]. To prepare the material for high wear applications, such as those required for ground engaging components, high chromium white iron will be infiltrated into the lattice matrix through liquid-solid bimetal casting techniques to form a bimetallic composite of high hardness, wear resistant white iron reinforced by an interior lattice of high strength steel [5,7].

The fusion zone between the two materials will be examined through microstructure studies to determine the integrity of the composite bond. In this manner, a bimetallic composite will be examined which is produced using a hybrid manufacturing process blending metal additive methods with conventional casting techniques.Alloy DevelopmentThe production of a bimetallic composite with a robust interface is heavily dependent on alloy development and selection. For this reason, the two alloys used in this composite should be carefully selected so that they exhibit the necessary individual material properties while also being highly compatible with each other. AM is the most heavily restricted manufacturing process in terms of alloy selection and thus the AM alloy will be selected first. As the casting process accepts a much wider range of alloy compositions, the cast alloy will be subsequently developed to compliment the choice of AM material.

As the casting process accepts a much wider range of alloy compositions, the cast alloy will be subsequently developed to compliment the choice of AM material. In this manner two alloys will be selected which promote complete fusion between the materials.Stainless steel (SS) 316L was selected as the lattice structure material for L-PBF due to its excellent additive manufacturability, corrosion resistance, low cost, and availability. This is a high alloy stainless steel which features up to 18 wt% Cr, which gives it excellent corrosion resistance necessary for ground engaging applications [33]. 316L stainless is additionally considered to be a weldable steel, featuring good resistance to carbide formation in the fusion zone [33].

316L stainless is additionally considered to be a weldable steel, featuring good resistance to carbide formation in the fusion zone [33]. The high chromium content of 316L also makes it a good match with high Cr white iron, as Cr diffusion745from the stainless steel will enhance the properties of the infiltration alloy rather than detract from them.High Cr white iron was selected as the infiltration alloy in this experiment; however, flexibility exists in the exact chemistry of the alloy. To determine the exact composition of white iron for this experiment, the effects of alloying elements on the final material properties were considered. From these recommendations a final composition was selected which satisfies the property requirements of the infiltration alloy.First, the infiltration abilities of the white iron alloy must be considered. White iron alloys are typically low in carbon compared to other cast irons, and therefore suffer from poor fluidity.

White iron alloys are typically low in carbon compared to other cast irons, and therefore suffer from poor fluidity. In order for infiltration to be successful, the fluidity of the metal must be improved. This can be accomplished through increased Si content, which improves the fluidity of the alloy [34]. Standard composition ranges for white iron are published by the American Society for Metals (ASM) and are heavily referred to in developing this alloy. High, medium, and low alloying element content is referred to in direct reference to the published content ranges for white iron. [34] Thus, to ensure that the fluidity of the white iron is such that the alloy can easily infiltrate throughout the lattice structure, Si content in the upper ranges of acceptable content was selected.Next, the Ni and carbon content of the alloy was identified to ensure optimal wear performance. Ideal wear performance in white iron is typically obtained by an iron matrix of heat- treated martensite [34].

Ideal wear performance in white iron is typically obtained by an iron matrix of heat- treated martensite [34]. For a successful heat treatment without cracking or distortion, it is preferred that the iron matrix be dominated by an as-cast austenitic structure. Because of this, Ni is typically included to inhibit pearlite formation [34]. If pearlite is properly suppressed, the iron matrix becomes super-saturated with carbon and chromium upon cooling and a fully austenitic matrix can be achieved. Although martensite is beneficial for wear resistance, martensite can be brittle and show poor toughness unless properly controlled. Lath martensite is a martensitic structure which features fine grain structure and better toughness than its counterpart, plate martensite.

Lath martensite is a martensitic structure which features fine grain structure and better toughness than its counterpart, plate martensite. Lower C content favors lath martensite formation and therefore C on the lower end of the acceptable range was selected [35].High Cr white iron is exceptionally good for abrasion resistance due to the formation of Cr carbide (Cr7C3 or Cr23C6). For this reason, high chromium content is essential. Additionally, a gradient between the Cr content of the white iron alloy and the 316L stainless alloy would be desired to foster diffusion across the interface, improving the bonding potential. However, it is important to consider the effects of the carbide formation within the interface of the two materials. Excessive carbon and Cr will diffuse to the grain boundary, leading to carbide precipitation and growth [36].

Excessive carbon and Cr will diffuse to the grain boundary, leading to carbide precipitation and growth [36]. Carbide formation in the grain boundaries can cause increased sensitization to Intergranular Stress Corrosion Cracking (IGSCC) due to Cr content depletion [37,38].Alloy ranges typical for high Cr white iron are shown in Table 1. These are obtained from standards published in the ASM Handbook for high alloyed cast irons [34]. Alloy ranges for carbon, silicon, nickel, chromium, and manganese specifically were identified.746Table 1: Recommended alloy ranges for high Cr white iron [34]Recommended Range (wt%)C 2.7-3.6Si 1.0-1.5Ni 3.4-5.5Cr 9-15Mn 0.8 maxBecause impact resistance is important in ground-engaging abrasion resistant components, carbon on the lower end of the acceptable spectrum was selected (2.7 wt% C). Higher Si than typical, 2.0 wt% Si, was chosen to improve the fluidity and ensure complete infiltration of the lattice.

Higher Si than typical, 2.0 wt% Si, was chosen to improve the fluidity and ensure complete infiltration of the lattice. 4.8 wt% Ni was chosen due to recommendations from the ASM handbook on pearlite inhibition at large casting thicknesses. Cr on the lower end of the acceptable spectrum was selected due to expected diffusion from the 316L high chromium stainless steel lattice. Lower Cr was additionally selected to drive diffusion across the interface between the high Cr in the stainless steel near 16 wt% and the lower Cr in the white iron near 10%. Lastly, a maximum of 0.8 wt% Mn was allowed as it is an austenite stabilizer and increases as cast austenite. The final selected composition of high Cr white iron is shown in Table 2.Table 2.

Lastly, a maximum of 0.8 wt% Mn was allowed as it is an austenite stabilizer and increases as cast austenite. The final selected composition of high Cr white iron is shown in Table 2.Table 2. Final selected composition for high Cr white iron [34]Recommended Range (wt%)C 2.7Si 2.0Ni 4.8Cr 10Mn 0.8Lattice DesignAs a tetrakaidecahedron unit cell structure has been identified in prior research as possessing superior strength compared to other unit cell morphologies, it was selected for use in this research [21]. The Lattice Commander module in Autodesk Netfabb was used to convert a 3D solid into a tetrakaidecahedron lattice structure. A relatively small unit cell size of 15mm was selected as the starting point for infiltration examination. A cube was selected as the lattice geometry for infiltration for ease of sectioning, polishing, and observing the face of the component for fusion between the two materials.

A cube was selected as the lattice geometry for infiltration for ease of sectioning, polishing, and observing the face of the component for fusion between the two materials. The cube size was selected to be 60 mm, a multiple of the unit cell size. A rendering of the 15mm unit cell lattice with a 1.5mm strut diameter is shown in Figure 1.747Figure 1: Tetrakaidecahedron lattice developed with Netfabb Lattice CommanderLattice FabricationThe lattice structure was printed using a Farsoon FS271M L-PBF System [39]. Supports were not included, as the bottom layer was considered sacrificial and would be cut off when the part was removed from the print bed. The machine was cleaned and then loaded with 316L stainless powder, consisting of a -45+15 particle size distribution and processed through a 200 mesh sieve. The chamber was sealed and preheated to 100°C prior to printing. The chamber was purged with argon to maintain a low oxygen content.

The chamber was sealed and preheated to 100°C prior to printing. The chamber was purged with argon to maintain a low oxygen content. General processing parameters used are shown in Table 3. The layer thickness selected provides adequate z-height feature resolution without significantly decreasing print time, while using common power and feeds with rotating hatch directions for SS316L. Multiple lattice structures were printed to ensure the success of the experiment. Upon completion of the print, parts were removed from the substrate using a horizontal bandsaw [39].Table 3. Processing Parameters used for Fabrication of SS316L Lattice Structures Processing Parameter Description Layer Thickness Laser Power Scan Speed Hatch DistanceValue 30 µm 225 W 1000 mm/s 0.090 µm748MAGMA Simulation SetupTo ensure optimum filling, MAGMA metal casting simulation software was used to predict the filling/infiltration behavior within the lattice structure.

The primary concern with fully infiltrating the lattice with white iron was the lattice acting as a chill, causing the liquid metal to freeze off or loose fluidity before full infiltration. To ensure that a gating design and pouring temperature were selected which would prevent this from occurring, a casting simulation was developed. The effectiveness of a side fill gating design and a top fill gating system design, shown in Figure 2, were explored using MAGMA. It was hypothesized that the side fill design would be better suited to promote smooth filling due to the low turbulence characteristics of the bottom up fill trajectory. Material parameters for high Cr white iron were selected from the MAGMA database for the pouring alloy, and readily available green sand was selected as the mold material.

Material parameters for high Cr white iron were selected from the MAGMA database for the pouring alloy, and readily available green sand was selected as the mold material. The internal lattice was set to be a steel chill with the material properties selected from an existing MAGMA database; the differences between the standard steel chill and a chill composted of SS316L were considered to be minimal.a)b)Figure 2. Lattice Infiltration Gating Designs: a) side fill and b) top fillPattern Design and ProductionFollowing the adoption of the optimal gating system, a pattern was designed and printed for preparing the mold. A pouring cup and square downsprue system was selected to minimize turbulence in the liquid and to ensure the pouring height remained consistent. A split pattern was designed which featured a 1.5° draft to all sides, allowing for easy mold release. Alignment features were also included for ease of mold making.

A split pattern was designed which featured a 1.5° draft to all sides, allowing for easy mold release. Alignment features were also included for ease of mold making. The gate was modeled based on MAGMA simulation results and was included in the drag side of the split pattern mold. The pattern was 3D printed using a Creality CR-10s Pro FFF printer out of PLA, using a 20% infill to reduce material consumption while maintaining adequate strength to resist compression during molding.Casting PreparationsPrior to lattice infiltration, a charge table was developed based on materials on hand at Georgia Southern University for the melting of the desired high Cr white iron. The charge table749was designed to account for the raw material chemistries and combined them in such a way that the desired chemistry was achieved. Each material in the table contains a specific weight percent of elemental values that, when combined together, give the desired chemistry of white iron.

Each material in the table contains a specific weight percent of elemental values that, when combined together, give the desired chemistry of white iron. The charge constituents and their quantities are shown in Table 4.Table 4: Charge table for selected high Cr white iron chemistryIn charge Pig Iron 1018 Steel FeSi FeMn FeCr Ni Total chargeWeight (kg) Weight (lbs) 11.5 2.43 0.45 0.12 2.6 0.86 17.9637.4 7.70 1.28 0.40 8.58 2.86 39.51The charge materials were carefully weighted out and added to a 60 lb induction furnace. Next, green sand was prepared for molding. Green sand was mulled and water was added until the compactability reached 45% compaction according to green sand testing equipment. The mold was then prepared as shown in Figure 3. Two cavities were poured to ensure a successful casting was obtained. The iron was poured at 1500 °C.

The mold was then prepared as shown in Figure 3. Two cavities were poured to ensure a successful casting was obtained. The iron was poured at 1500 °C. It was first de-slagged and then transferred to a secondary ladle for pouring.DragCopeDownspruea)Gateb)Figure 3: Greensand a) cope and b) drag used to infiltrate lattice with high Cr white ironPost ProcessingAfter casting, the mold was shaken out. The gating system was removed using a reinforced cut-off wheel mounted on an angle grinder. The casting was sectioned using Wire EDM, as the hardness of the material prevented traditional sectioning methods from being used. A square section showing the interface between the unit cell and the infiltration allow was obtained and mounted as shown in Figure 4. The bottom surface of the sample was ground flat and polished to a 3 μm finish for microscopy. The surface was deep etched with Marbles Reagent to bring out the750carbide phase for imaging.

The bottom surface of the sample was ground flat and polished to a 3 μm finish for microscopy. The surface was deep etched with Marbles Reagent to bring out the750carbide phase for imaging. Metallography was then performed on an Olympus BX 53M optical microscope at varying magnifications to examine the diffusion interaction between the materials.High Cr white ironStainless Steel 316LFigure 4: Sample Prepared for Microstructure AnalysisSimulation ResultsFrom the MAGMA simulation results shown in Figure 5, it is clear that the side fill gating system is the more effective means of filling when compared to top fill gating. The top fill method resulted in high turbulence due to the rain-fall type filling around the lattice which drastically increases the risk of defects, oxide formation, and poor bonding. The side fill gating system minimized the overall flow turbulence and promoting smooth filling of the lattice structure.

The side fill gating system minimized the overall flow turbulence and promoting smooth filling of the lattice structure. For this reason, the side fill gating system was chosen for the final casting system.Figure 5: MAGMA Simulation Results Comparing the Two Gating System Designs751Additive Manufactured Lattice ResultsThe lattice structures printed via L-PBF are shown in Figure 6. The overall print quality of the cell lattice structures appeared to be high, with few defects observed. The surface finish was uniform, although some excessive roughness did occur at the overhang sections. Very few other defects existed, with most occurring at the end of the print due to cumulative damage incurred on the recoater. However, for the purpose of infiltration and fusion studies, all structures were adequate for study.Figure 6: 316L Stainless Lattice Surface FinishCasting Quality and Infiltration ResultsA final casting after the gating system was removed is shown in Figure 7.

The samples were oxidized (scale & ferrous oxide) and had remnants of green sand burn-on. The outside surface finish was not measured but appeared to be slightly rougher due to high pouring temperature. For typical high temperature steel castings, a zircon mold wash is applied to prevent sand-metal interactions, although this is not typically necessary for cast iron. However, since white iron features a lower carbon content than typical cast iron and thus requires a higher pouring temperature, its interaction with the greensand mold material was similar to that of steel castings, and a mold wash would have been beneficial to improve the surface finish. As casting surface quality is not critical to the success of the infiltration experiment, the castings were overall deemed to be successful.

As casting surface quality is not critical to the success of the infiltration experiment, the castings were overall deemed to be successful. Scale and ferrous oxide were removed from the areas of interest using sand blasting and grinding before sectioning for microscopy.752Figure 7: As-Cast 316L Stainless Steel Lattice Infiltrated with high Cr white ironChemistry AnalysisSamples taken from the gating systems of the castings were used for elemental analysis through optical spectroscopy final as-cast chemistry. An Amtek the SPECTROMAXx spectrometer calibrated based on “type correction” standards was used to burn the samples and obtain the actual chemistry. Three burns were made, and the results were averaged.

Three burns were made, and the results were averaged. The resulting chemistry according to the spectrometer reading is shown in Table 5 and compared with the target chemistry.to determineTable 5: Target versus Actual CompositionTARGET COMPOSITION (WT%) 2.7 C 2.0 SI 4.8 NI 10 CR MN 0.8ACTUAL COMPOSITION (WT%) 2.8 2.11 4.25 9.84 0.82PERCENT ERROR (%)3.7 5.5 11.5 1.6 2.5Using this measured chemistry, a solidification diagram was generated using the JMat Pro Chemistry Analysis software and is shown in Figure 8. The solidification diagram gives information as to what phases will be present in the final as-cast microstructure.

The solidification diagram gives information as to what phases will be present in the final as-cast microstructure. As seen in the diagram, the final microstructure will feature an austenitic matrix with carbide and cementite phases.753Figure 8: Solidification diagram generated for the actual white iron chemistryMicrostructure AnalysisFollowing the polishing and etching of the bimetal composite interface, metallographic images were taken at varying magnifications.

The high Cr white iron microstructure was examined and is shown in Figure 9.Figure 9: Medium magnification image of the white iron microstructureMedium and high magnification images of the fusion zone between the stainless steellattice and high Cr white iron are shown in Figure 10.754Stainless Steel Latticea)b)Bimetal Fusion ZoneHigh Cr white ironFigure 10: a) Medium and b) high magnification image of carbide growth along the grain boundary in the bimetal fusion zoneIn examining the microstructure of the high Cr white iron, extensive carbide formation is observed in the white iron. Figure 9 resolves the Cr carbide network through an austenitic matrix. This is expected based on the solidification diagram generated by JMat Pro thermal analysis software shown in Figure 8. Additionally, the microstructure shown in Figure 10(b) at the bimetal interface shows carbide precipitating out along the grain boundary of the stainless steel austenitic microstructure.

Additionally, the microstructure shown in Figure 10(b) at the bimetal interface shows carbide precipitating out along the grain boundary of the stainless steel austenitic microstructure. This grain boundary carbide precipitation indicates diffusion along the interface between the two alloys.Figure 10 also shows an example of complete fusion between the AM lattice structure and the cast white iron. As can be seen, complete metallurgical bonding was achieved with little if any porosity, inclusions, or other defects. This was observed throughout the interior of the cast samples.Although throughout the interior of the structure complete metallurgical bonding was observed, incomplete fusion did occur at the outer walls of the sample. Along the outer bounds of the bimetallic structure, at the interface between the two alloys, complete fusion was not observed.

Along the outer bounds of the bimetallic structure, at the interface between the two alloys, complete fusion was not observed. Distinct separation of the lattice and high Cr white iron matrix did occur in some places, as shown in Figure 11.755Figure 11. Separation of the stainless steel lattice and high Cr white iron along the outer bounds of the structureThe incomplete bonding observed at the surface of the sample is likely due to contact with the mold wall. Effective bonding between the two alloys is the function of heat and time. Due to heat transfer away from the structure to the mold wall, the interior of the sample had more time for fusion at temperature as compared to the outer surface.

Due to heat transfer away from the structure to the mold wall, the interior of the sample had more time for fusion at temperature as compared to the outer surface. Although incomplete fusion was observed at the surface of the sample, complete infiltration of the lattice at these locations resulted in a firm mechanical fit between the two materials.Through this casting process, a bimetallic composite structure was achieved through both metallurgical and mechanical bonding of two materials. Through the implementation of an AM lattice structure, brittle white iron was reinforced by a ductile stainless steel lattice structure, successfully resulting in a reinforced bimetallic wear component.Future WorkIn order to further quantify the effects of integrating stainless steel 316L with high Cr white iron, the authors propose to investigate both mechanical and microstructural properties.

EDS analysis across the fusion zone will enable a full gradient of Cr depletion and carbide formation along the grain boundary. Previous studies have shown that carbide precipitation is less for large grain microstructures, and greater for small grain microstructures [38], therefore, characterization of grain size will qualitatively support the EDS analysis. Furthermore, micro hardness testing will provide a gradient of micro-scale mechanical properties across the fusion zone while fatigue and quasi-static testing will enable macro-scale properties. Simulated mechanics models will be compared against the experimental results and adjusted accordingly.

Simulated mechanics models will be compared against the experimental results and adjusted accordingly. Additionally, the starting chemistry and molding techniques can be altered in light of the current results to promote diffusion and control carbide morphology, leading to solid solution strengthening.ConclusionsA method has been introduced in which high abrasion resistant alloys such as high Cr white iron may be reinforced through the application of AM techniques. It is shown that a stainless steel lattice structure produced through L-PBF may be infiltrated with high Cr white iron to produce a756reinforced bimetallic wear component. It was observed that complete lattice infiltration was possible, and a mixed bimetal interface was observed featuring fused and unfused interface zones. The metallurgically bonded interface region featured chromium diffusion from the AM lattice structure which formed chromium carbides along the interface between the two materials.

The metallurgically bonded interface region featured chromium diffusion from the AM lattice structure which formed chromium carbides along the interface between the two materials. This research shows that high Cr white iron may be used to infiltrate an additively manufactured stainless steel lattice structure for producing a reinforced wear component.757References[1]El-Aziz, K. A., and Saber, D., 2020, “Mechanical and Microstructure Characteristics of Heat-Treated of High-Cr WI and AISI4140 Steel Bimetal Beams,” J. Mater. Res. Technol., 9(4), pp. 7926–7936.[2] Wróbel, T., 2014, “Characterization of Bimetallic Castings with an Austenitic Working Surface Layer and an Unalloyed Cast Steel Base,” J. Mater. Eng. Perform., 23(5), pp. 1711– 1717.[3] Rong, S. F., Zhao, X. M., Zhou, H. T., Wang, J. F., Yao, J., and Li, C. H., 2014, “Study on Coated-Liquid-Liquid Trimetal Composite Casting Hammer,” Mater. Res. Innov., 18(5), pp.

J. Fatigue, 94, pp. 211–217.[31] Anand, M., and Das, A. K., 2021, “Issues in Fabrication of 3D Components through DMLSTechnique : A Review,” Opt. Laser Technol., 139(February 2020), p. 106914.[32] Simchi, A., and Asgharzadeh, H., 2004, “Densification and Microstructural Evaluation during Laser Sintering of M2 High Speed Steel Powder,” Mater. Sci. Technol., 20(11), pp. 1462–1468.[33] “316/316L Stainless Steel.” [34] Gundlach, R. B., 2018, “White Iron and High-Alloyed Iron Castings,” Casting, 200(c), pp.896–903.[35] Bhadeshia, H. K. D. ., and Honeycombe, R. W. ., 2017, Steels: Microstructure and759Properties, Elsevier, Cambridge.[36] Cao, T., Cheng, C., Ye, F., Xv, H., and Zhao, J., 2020, “Relationship between Carbon Segregation and the Carbide Precipitation along Grain Boundary Based on the Structural Unit Model,” J. Mater. Sci., pp. 7883–7893.[37] Yin, Y., Faulkner, R.

nder Development for Indirect SLS of Non Metallics Kumaran M. Chakravarthy and David L. Bourell Laboratory of Freeform Fabrication, The University of Texas at Austin, Austin, TX-78712Abstract: In this work, a binder system for non metallics such as graphite has been developed and tested. A graphite-phenolic mixture was used for making bipolar plates using SLS. Phenolic resin melts (~90°C) instead of cross-linking (~165°C) in the SLS machine. During post processing in a furnace, the SLS parts slumped due to remelting of phenolic resin with no other binder present to hold the powder together. To prevent this slumping and to increase the green strength of the SLS parts, a new binder system was developed with Nylon/11 added. Nylon/11 has a melting point of ~185°C and is able to hold the SLS part together until the phenolic can cross link, preventing slumping.

Nylon/11 has a melting point of ~185°C and is able to hold the SLS part together until the phenolic can cross link, preventing slumping. Single layer scan experiments with different compositions and SLS parameters were run to identify optimum composition and SLS parameters. Three point bend specimens were prepared with 70 wt. % Graphite-25 wt. % Phenolic- 5 wt. % Nylon/11 and 70 wt. % Graphite-20 wt. % Phenolic- 10 wt. % Nylon/11, tested and compared to graphite-phenolic parts. No significant increase in green strength was attained with Nylon/11 addition.1. Introduction: In indirect selective laser sintering (SLS) high melting temperature materials like metals and ceramics are coated or mixed with polymers and are then processed in the SLS machine. Under the laser power, these polymers melt and bind the inorganic materials together.

Under the laser power, these polymers melt and bind the inorganic materials together. Successful use of metals and ceramics in SLS process depends on successful development of binders.Commercially, there are few material systems such as LaserformTM A6 steels available which are tool steels coated with unspecified binders [1]. These binders serve as sacrificial binders in that they are useful only in the initial stages: processing in the SLS machine and to the start of furnace sintering. Once sintering starts, the component/part gets strength from sintering and these binders are eliminated by high temperature burnout. There are some applications wherein sintering is difficult as in the case of graphite/SiC and a binder that can give strength during all the stages of processing either by remaining as itself or by transforming to some other state is needed.

One such application is the fabrication of graphite bipolar plates or current collectors in fuel cells.Research has been done on development of binders for graphite bipolar plates [2] and SiC composites and wafer carrier plates [3, 4]. Chen [2] used Phenolic resin as a binder. The Phenolic resin was pre-mixed with a hardener/ catalyst and was purchased from Georgia Pacific. Novalac was the resin and had HMTA or hexamine was the catalyst. The green strength of the Graphite-Phenolic parts was not good. The green strength values were not reported. After infiltrating with some resin such as epoxy after furnace curing, the parts gained strength. Bhandari [3] tried to improve the green strength of SiC parts by adding Nylon and Wax in addition to Phenolic resin. Trials/experiments were done outside theSLS machine by mixing powders, making three point bend specimens and heating them in a furnace at ~140°C. Green strength in the range of ~7 -19 MPa was achieved.

Green strength in the range of ~7 -19 MPa was achieved. It is difficult to get such a long exposure to temperature in a SLS machine. A value of 2 MPa will ensure that the parts have adequate strength to be handled.Another motivation in identifying a suitable binder composition is that Phenolic melts (~90° C) before it can cross link (~165°C) as seen in the DSC plot below [2].Figure 1: DSC plot of Phenolic resin [2] shows the resin melts much (50-90°C) before cross linking occurs (~150-160°C)During post processing in the furnace, the SLS parts slump due to remelting of the phenolic resin with no other binder to hold the powder together. To prevent this slumping and improve the green strength, its necessary to identify some binder composition that can hold the powder together in the range of 90-165°C. Nylon 11 has a melting point of ~185°C, so it can be used to retain the shape in that temperature range.2.

Nylon 11 has a melting point of ~185°C, so it can be used to retain the shape in that temperature range.2. Experimental:The graphite-Phenolic resin- Nylon/11 were mixed with different composition and heated in a oven to ~ 200°C. Based on the nature of the mix (how hard or soft they were), specific compositions were identified to be suitable for further processing. These compositions were then used in single layer scan experiments in the SLS machine DTM sinterstation 2000 to identify optimum process parameters. Based on these results, 3 point bend specimens were prepared.2.1 Materials: Graphite (GS 150 E) was purchased from Graftech International. Phenolic resin (Durite AD 332 A) was purchased from Hexion in a premixed condition.

Phenolic resin (Durite AD 332 A) was purchased from Hexion in a premixed condition. Nylon/11 was purchased from Rilsan Fine Powders.Graphite- Phenolic-Nylon were mixed in the following ratio, placed in a oven and heated to ~200°C to observe the nature of the mix:Graphite (wt%) Phenolic (wt%) Nylon11 (wt%)Remarks70300Base composition70255702010701515Table 1: Initial compositions used in oven testOf these, 5 & 10% Nylon addition made the mass harder than the base composition and 15% Nylon addition made the mass softer than the base composition. Based on these experiments, we identified that 5% & 10% Nylon 11 addition as two potential compositions.2.2 Single layer scan experiments:To confirm the behavior of these mixes in the SLS machine under laser power and to minimize the use of powder, single layer scan experiments were performed with different laser power to identify optimum laser parameters.

Single layer scan experiments were done in a DTM Sinterstation 2000. The build height was kept at 0.003 inches and the powder layer thickness was kept at 0.004 inches to prevent the rollers from moving. Graphite powder was mixed with different binder composition and laid on a graphite plate on the partbed. In all the experiments, the weight ratio of Graphite to binder was kept at 70:30. This composition was identified by Chen [2]. The compositions used for single scan experiments were same as shown in table 1.SLS parameters used: For Single layer scan experiments, the partbed temperature was kept at 80°C and the feed bin temperature was kept at 45°C. Fill Laser powers used were 7, 12, 17 W (12 W was identified to be optimum by Kaushik [5]) and an outline laser power of 4 W was used.Part bed for single layer scan exptA single layer scan in progressFigure 2: Single layer scan experiment in DTM sinterstation 2000.Different shapes were created using different laser power.

Ability to produce sharp features, relative green strength to handle the single layer scan parts and absence of curling or delamination were the metric used to identify the suitability of a particular binder composition.2.3 Three point bend specimens by SLS process: 2.3.1 5 wt. % Nylon: 2.3.1.1 Powder preparation: Based on the results from single scan experiments, the 5% Nylon (Graphite: 70 wt% - Phenolic: 25 wt%- Nylon: 5%) mix was processed on a scale enough for a build in the SLS machine. The powders were mixed in the above proportion in a ball mill with alumina media for 24h. The powders were sifted to remove coarse particles and were filled in the SLS machine.2.3.1.2 SLS Build: Three point bend specimens (5 in X 0.5 in X 0.125 in) were prepared using 12 and 14 watts laser power (identified by single scan experiments). Samples did not have acceptable strength and broke during removal from the partbed. A new build was made using the following laser powers: 18 W, 22 W, and 26 W.

Samples did not have acceptable strength and broke during removal from the partbed. A new build was made using the following laser powers: 18 W, 22 W, and 26 W. Four or five specimens were made with each laser power. Two builds were made to assess the repeatability.2.3.2 10 wt. % Nylon: 2.3.2.1 Powder preparation: The 5% mix prepared for previous builds was modified to 10% Nylon composition by adding extra Nylon/11 & graphite. The 5% mix was sifted, new additions (extra graphite and Nylon/11) were made and the mix was ball milled for 24 h. The mix was again sifted to remove any large lumps of powder.2.3.2.2 SLS build: Three point bend specimens (5 in X 0.5 in X 0.125 in) were prepared using 22, 26, 28 and 30 watts laser power. 22 & 26 W were used in the 5 wt. % Nylon experiments. 28 & 30 W were used to see if increasing the laser power had an effect on the strength. The partbed temperature was kept at 80°C and the feed bin temperature was kept at 45°C.

% Nylon experiments. 28 & 30 W were used to see if increasing the laser power had an effect on the strength. The partbed temperature was kept at 80°C and the feed bin temperature was kept at 45°C. The outline laser power was 4 watts.Four or five specimens were made for each laser power in each build. Three builds were made in all to assess the repeatability.2.4 Mechanical characterization: The specimens were tested in a 5 kN Instron machine (Model 3345) with three point bend setup as shown in Figure 3.Figure 3: A three point bend test setup.The span length was set at 50 mm and a displacement rate of 0.1 mm/min used. A Flexure stress versus flexure strain plot was obtained in all the tests and the maximum flexure stress was reported as the flexure strength.3. Results:3.1 Single layer scan experiments:Figure 4 below shows the single scan layers (the numbers in the figures denote the wt.

Results:3.1 Single layer scan experiments:Figure 4 below shows the single scan layers (the numbers in the figures denote the wt. % of Graphite, Phenolic, nylon/11 respectively).70-30-070-25-570-22.5-7.570-20-1070-15-15Figure 4: Single layer scan parts with different binder compositionAll the binder compositions were able to produce the features and none had any curling or delamination. The dimension of the parts were measured and compared to the dimensions of the input drawing. 15% Nylon showed more shrinkage than the rest. The strength of nylon- added parts was relatively better than graphite-phenolic parts. Some graphite-phenolic parts broke while trying to remove them from the partbed, while the ylon- added parts did not break. This showed that there was some relative improvement n due to Nylon addition.B % Nylon additions.ased on these results, 3 point bend specimens were built with 5 wt. % Nylon and 10 wt.3 .2 Three point bend test results: 3.2.1 5 wt.

% Nylon and 10 wt.3 .2 Three point bend test results: 3.2.1 5 wt. % Nylon mix:Stress - strain plots of 5% Nylon mix with different laser power are shown in the Figures 5 and 6 below:f Figure 5: A typical 5% Nylon mix with 22 W ill laser power. The flexure strength of this specimen is 1.42 MPa.Figure 6: A typical 5% Nylon mix with 26 W fill laser power. The flexure strength of this specimen is 1.54 MPa.It was found that 5% Nylon addition did not improve the green strength dramatically. The green strength of compacts used in power metallurgy is around 7 MPa [6]. This 7 MPa green strength enables the compact to be handled. A realistic target of 4-5 MPa was set which should be sufficient enough to excavate the parts from the partbed and transfer them to the furnace.3.2.2 10 wt. % Nylon addition The flexural stress-flexural strain plots for different laser powers are shown in Figure 7 (22 watts), Figure 8 (26 watts), Figure 9 (28 watts) and Figure 10 (30 watts).

% Nylon addition The flexural stress-flexural strain plots for different laser powers are shown in Figure 7 (22 watts), Figure 8 (26 watts), Figure 9 (28 watts) and Figure 10 (30 watts). The associated tables list point values. (Note: the difference in the shape of the curve was due to the different pre-loading used before the start of the test).Figure 7: Flexural stress-strain plots of 3 point bend specimens made with 22 watts laser power.Table 2: Flexural load and stress values of 3 point bend specimens made using 22 Watts laser power & 10 wt. % NylonFigure 8: Flexural stress-strain plots of 3 point bend specimens made with 26watts laser power.Table 3: Flexural load and stress values of 3 point bend specimens made using 26 Watts laser power & 10 wt. % NylonFigure 9: Flexural stress-strain plots of 3 point bend specimens made with 28watts laser power.Table 4: Flexural load and stress values of 3 point bend specimens made using 22 Watts laser power & 10 wt.

% NylonFigure 10: Flexural stress-strain plots of 3 point bend specimens made with 30 watts laser power.Table 5: Flexural load and stress values of 3 point bend specimens made using 30 Watts laser power & 10 wt. % NylonThe flexural strength of 10 wt. % Nylon specimens was ~1 MPa. Increasing the Nylon/11 content degrades the strength values.3.3 Carbonization experimentsTo observe the slumping behavior of the green parts, four nylon added specimen were put in a high temperature vacuum furnace along with two graphite-phenolic parts. They were heated to 1200°C. All the parts slumped as shown in figure 11. Nylon/11 additions didn’t help in preventing the slumping.Figure 11: 3 point bend specimens in the furnace before and after vacuum furnace processing. All the parts were slumped after carbonization at 1200°C.Some carbonized samples were tested for the brown flexural strength. All the brown (carbonized) parts showed an increase in flexural strength (~2.5-2.75 MPa).

All the brown (carbonized) parts showed an increase in flexural strength (~2.5-2.75 MPa). A typical flexure stress- flexure strain plot of the brown parts is given below:Figure 12: A typical flexure stress- flexure strain curve for a brown (Carbonized) part. The flexural strength of the brown parts was in the range of ~2.5-2.75 MPa.4. Discussion:Both the 5 wt. % and 10 wt. % Nylon specimens had a flexural strength of ~ 1-2 MPa. Higher addition of Nylon/11 apparently degrades the strength. This value is comparable to the green strength value of LaserFormTM Tool steel of 3D systems. LaserFormTM has Phenolic, Nylon and wax in its binder composition and has a green strength of 1-2.1 MPa. This could be the maximum possible strength value with this composition.

LaserFormTM has Phenolic, Nylon and wax in its binder composition and has a green strength of 1-2.1 MPa. This could be the maximum possible strength value with this composition. Higher strength could possibly be achieved only by a significant change/ new additions to the binder composition.Also, the particle size of the Phenolic resin could be another factor that influenced the green strength adversely. The earlier experiments [2, 5] used cryomilled phenolic resin, which had particle sizes in the range of ~10-15 microns. The phenolic resin used in this experiment was unmilled with particle size in the range of ~31 microns. So, the phenolic particle size would have nullified the nylon effect.

The phenolic resin used in this experiment was unmilled with particle size in the range of ~31 microns. So, the phenolic particle size would have nullified the nylon effect. Scanning electron microscope images of some of these specimens were obtained to see if there was any difference between the milled and un- milled resins’ binding behavior.The SEM images of a specimen prepared using milled phenolic powder are shown in the Figure 13 below:Figure 13: SEM images of a specimen prepared using milled phenolic resin. Note the dispersion of the resin is uniform and no big globs of molten resin seen.The SEM images of a specimen prepared using un-milled phenolic powder are shown in the Figure 14 below:Figure 14: SEM images of a specimen prepared using milled phenolic resin. Note the large globules of molten resin seen, marked by red circles.The milled resin sample did not have any large molten globules.

Note the large globules of molten resin seen, marked by red circles.The milled resin sample did not have any large molten globules. But the un-milled resin sample had large molten globules indicating that the dispersion behavior was different.5. Conclusion: It appears that Nylon addition up to 10% does not provide improvement in the green strength of the SLS graphite parts. Nylon/11 addition also does not prevent slumping. One possible reason for this behavior (effect on strength) could be due to the initial particle size of the powder. Phenolic resin used by Chen [2] & Kaushik [5] was cryogenically milled to 10-15 micron size whereas the resin used in this work is unmilled and was around 31 microns. Particle size has an effect on the sintering behavior and so on the green strength. Reducing the phenolic resin particle size, addition of acrawax resin to the composition and efforts to induce cross-linking inside the SLS machine are future plans to improve the green strength.

Reducing the phenolic resin particle size, addition of acrawax resin to the composition and efforts to induce cross-linking inside the SLS machine are future plans to improve the green strength. Acknowledgement: This work was supported by the Office of Naval Research MURI grant no. N00014-07-1- 0758. The authors thank the support of Hexion Specialty chemicals Inc., Louisville, KY in supplying Phenolic resin used in this work.References:1. LaserformTM A6 Steel Material, Materials guide 3D system. 2. Fabrication of PEM Bipolar plates by Indirect selective laser sintering, SS Chen, PhD Dissertaion, The University of Texas at Austin, 20063. Binder optimization for manufacturing of Silicon carbide performs, S Bhandari, MS Thesis, The University of Texas at Austin, 20044. Indirect rapid manufacturing of Silicon carbide composites, RS Evans, PhD Dissertation, The university of Texas at Austin, 20055.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperBinder evaporation during powder sheet Additive Manufacturing J. Volpp1, W. Zhang2, W. Abbott2, A. Coban3, S. McConnell2, S. Marola4, R. Casati4, R. Padamati3, and R. Lupoi21 Department of Engineering Sciences and Mathematics, Luleå University of Technology, 971 87 Luleå, Sweden 2 Trinity College Dublin, The University of Dublin, Department of Mechanical, Manufacturing & Biomedical Engineering, Dublin, Ireland 3 School of Chemistry, CRANN, Trinity College Dublin, Dublin 2, Ireland 4 Department of Mechanical Engineering, Politecnico di Milano, Via G. La Masa 1, 20156, Milano, ItalyAbstractSeveral Additive Manufacturing methods are well established and found access into regular production in multiple sectors. For processing metals, typically wire or powder is used as feedstock.

For processing metals, typically wire or powder is used as feedstock. Wire processing is typically used for comparably large structure building, while powder processes offer, in general, a more precise metal application. For Powder Bed Fusion processes, very fine powder is used (typical 20 µm to 65 µm), while for Directed Energy Deposition powders are in the range between 50 µm and 160 µm. Such fine powders can be a health risk for humans (aspiration, skin integration). Avoiding contact with the powders in a production environment can be a big effort or not avoidable. Therefore, an alternative process was developed that provides the powder not as free powder particles but in form of powder sheets. For enabling the necessary bonding between the particles, a binder is used. In order to understand the impact of the binder during laser processing of the powder sheets, single pulse and line treatments were produced and recorded with high-speed imaging.

In order to understand the impact of the binder during laser processing of the powder sheets, single pulse and line treatments were produced and recorded with high-speed imaging. Recordings show the vaporization of the binder and the related ejections of powder particles. At lower energy input, the binder evaporation led to less spattering, which indicates that a binder heating at low heating rates induces less pressure on the powder particles.Keywords: Powder sheet processing, laser materials processing, vaporization, spatterIntroductionLaser powder Additive Manufacturing (AM) is an establish technology used by many actors in industry. Most established is the Laser Powder Bed Fusion with a laser beam of metals (PBF-LB/M), where a laser melts pre-placed powder layer by layer to produce complex structures (e.g. a combustion chamber [1] ).

a combustion chamber [1] ). Another technology is the Directed Energy Deposition with a laser beam of metals (DED-LB/M), where the powder is applied through a powder nozzle with gas assist, while the laser beam produces the melt pool, in which the powder particles incorporate (e.g. [2], [3]). Also wire-based processes are established, which have the advantage of high build-up rates at minimum material loss but are typically more useful for larger structures (e.g. [4]).Powders for laser AM are typically in the size range of 20 µm and 65 µm for PBF-LB/M and 50 µm and 160 µm for DED-LB/M. Such fine powders are efficient to form the powder layers or to be fed through the feeding systems but denote a health risk for humans during handling or exposure. Aspiration of powders can lead to lung diseases and fine powder particles can even access the human body through skin pores.

Aspiration of powders can lead to lung diseases and fine powder particles can even access the human body through skin pores. In particular in production facilities, it is required to make the necessary precautions for human safety and minimize human exposure to powder particles. However, handling requires human interaction and powder transfer in the air cannot be completely avoided. In order to reduce the exposure of metal particles to humans working in production facilities, a new approach is suggested in this work that promises limit extraction of loose powder.Therefore, a new methodology was successfully tested, called MAPS (Metal Additive manufacturing using Powder Sheets) [5]. This approach includes the pre-production of powder sheets. Regular powder material of nearly any powder size distribution and shape is mixed with a polymer binder and formed with a metal- polymer-solvent mixed solution, which is poured on a flexible Teflon sheet.

This is followed by the solvent evaporation from the mixture and the metal powder in conjunction with the polymer binder to solidify to form a flexible powder sheet. It is expected that the additional steps to incorporate powder particles into sheets will not324significantly increase the powder prices, but significantly increase the production worker´s health. Processing with a laser beam in a similar way to PBF-LB/M processing allows to form layer-by-layer structures.However, the process is not fully understood. In particular, the effect of the binder on the processing and metal transfer are unknown. Therefore, this work uses high-speed imaging in order to observe the laser interaction with the powder sheet in order to identify process characteristics.MethodologyPowder sheets were produced using 304L stainless steel powder (Carpenter Additive, powder size of 15-45 μm) forming powder sheets of 137 µm thickness (Fig. 1b).

1b). From the powder sheet production, the lower sheet side shows a flat binder/polymer surface (here called binder side) and the upper side, which is rougher, dominated by the powder (here called powder side). In order to process the powder sheets, a test setup (Fig.1a) was established using a scanner optics (Jenoptic) leading the laser beam from an IPG fiber laser (300W max power) to the free hanging powder sheet. The process zone was shielded by an Argon gas flow (18 L/min).a) Figure 1. a) Setup of the testing equipment and b) microscopic pictures of a powder sheetb)For observation of the laser impact, a high-speed camera (Photron) including illumination (Cavitar) was installed. The camera was positioned in horizontal position for the pulse experiments and in a 45° angle for the processing experiments (Fig. 1a). The impact of processing parameters was analyzed by varying the laser power and processing speed.

1a). The impact of processing parameters was analyzed by varying the laser power and processing speed. In addition, the processed powder side was varied.Powder sheets, before and after laser-material interaction, were characterized by means of a Zeiss Sigma 500 field emission scanning electron microscope (FE-SEM) equipped with an Oxford Instruments Ultim Max detector for energy dispersive X-ray (EDS) analysis.Results and DiscussionFirst, laser pulsing experiments were conducted by applying a 0.5 s pulse on different sides of the powder sheet. A pulsating vapor outflow to the top side was observed with a few small hot particle ejections. On the bottom side of the powder sheet, a vapor channel was established, which transports small powder particles and spatters in a more divergent way compared to the vapor stream (Fig. 2). Those tests show that a metal transfer should be possible with the ejected material, while the binder material vaporizes on the sheet top side.325Figure 2.

2). Those tests show that a metal transfer should be possible with the ejected material, while the binder material vaporizes on the sheet top side.325Figure 2. Representative high-speed frames of a pulsed laser impact on the powder sheet (horizontal camera position) and photograph of the powder sheet after processingMicrographs acquired on the area subjected to laser pulsing experiments showed the formation of some deposits enriched in C and O on the bottom side of the powder sheet. These deposits are present in a region approximately 300 µm around the hole left by the laser pulse. The investigation about the exact composition of the layer is ongoing. Several samples gave similar results regardless of the processed powder sheet side. The origin of such deposits is under current investigations.Figure 3. Representative micrographs of the sheet binder and powder side after the laser pulse. The red cross and dot on the micrographs indicate the direction of the laser beam.

Representative micrographs of the sheet binder and powder side after the laser pulse. The red cross and dot on the micrographs indicate the direction of the laser beam. The cross marks the side on which the laser beam was impinging the sheet, while the dot indicates the side opposite to the laser impact.When using line scanning on the free hanging powder sheets, different vapor and particle ejection behavior was observed (Fig. 4). When the powder side is on top, the vapor and spattering are more pronounced, indicating that the accumulated binder in the bottom part of the sheet expands and pushes powder particles upwards. In general, the vapor appearance is reduced at higher processing speeds as expected due to the lower energy input per time.

In general, the vapor appearance is reduced at higher processing speeds as expected due to the lower energy input per time. However, the sheets were not completely molten, and no material was ejected to the bottom at too high speed, which indicates that the processing window has a limit towards energy input that needs to be reached to achieve the material transfer and an upper limit where extensive vaporization disturbs the processing due to spattering.326Figure 4. High-speed images and powder sheet appearances at different process conditions at 200W laser powerWhen processing the powder sheets with a base sheet underneath, the process was observed (Fig. 5). Since the vaporized binder cannot escape downwards, a more pronounced vapor upwards is established compared to the free hanging sheet results. More extensive spattering occurs at higher laser power. In general, it was observed that less spattering and vapor occurs at low energy input.

More extensive spattering occurs at higher laser power. In general, it was observed that less spattering and vapor occurs at low energy input. At high laser power, material could be transferred to the base sheet forming an additively produced track.Figure 5. High-speed images, powder sheet and applied material appearances at different processing conditions304 stainless steel samples were printed by using the fabricated powder sheet and a commercialized Realizer SLM50 system in the AMBER AR-lab in TCD (Figure 6). 70 layers of 50×6×1.5 mm3 samples were printed, which shows the capabilities of the MAPS for printing large-scale structures. First visual inspection shows that after polishing, the cross section show nearly fully dense sample was printed (Figure 6b), comparable to that of the loose powder based PBF-LB/M. Quantitative and volumetric analysis is ongoing and will be presented in future work.327Figure 6.

Quantitative and volumetric analysis is ongoing and will be presented in future work.327Figure 6. 304 stainless steel sample printed by using powder sheet: a) three 50×6×1.5 mm3 samples and b) cross section of polished sample (5×5×0.6 mm3)Conclusions The experimental tests of the new Additive Manufacturing technology MAPS (Metal Additive manufacturing using Powder Sheets) show the feasibility of transferring material from the powder sheets to a base material. Observations of the vapor plume in high-speed images show the binder evaporation and related spattering. At lower energy input, the binder evaporation led to less spattering, which indicates that a binder heating at low heating rates induces less pressure on the powder particles.Acknowledgements The authors acknowledge the funding of PoSAddive – Powder Sheet Additive Manufacturing (EIT raw materials, No.

nder Formulation in EVA-wax system for Fused Deposition of CeramicsSenol Pekin * '1', Avigdor Zangvil *, and William Ellingson 'I' * Materials Research Laboratory, University of lllinois at Urbana-Champaign, Urbana, lL. 'I' Energy Technology Division, Argonne National Laboratory, Argonne, lL.Blends in the ethylene vinyl acetate (EVA) - wax system have been evaluated as potential binders.to be used in fused deposition of ceramics (FDC). In order to obtain good handling strength, it was indicated that the melting point of the polymer needs to be lower than that of the wax. In this context, it was shown that the melting point of the EVA decreases as the vinyl acetate content in the copolymer increases.

In this context, it was shown that the melting point of the EVA decreases as the vinyl acetate content in the copolymer increases. By measuring the viscosity as a macroscopic property, it was shown that 20 % vinyl acetate-containing EVA is miscible in microcrystalline wax up to, atleast, 30 %.Binders used in FDC need to have low viscosity and it was pointed out that the slumping can be one main problem associated with binders with low viscosity. Thus, thermolysis of wax at low temperature is suggested as a solution. The volumetric thermal expansion and melt strength of an EVA-wax blend were displayed as a function of temperature, in the form of a penetration test.I. IntroductionFused deposition of ceramics (FDC) has recently emerged as a new way of shaping ceramic components. [1]. The method has advantages in the sense. that it minimizes the product development time and cost; however, it also imposes several challenges regarding the ceramic processing.

[1]. The method has advantages in the sense. that it minimizes the product development time and cost; however, it also imposes several challenges regarding the ceramic processing. The main challenge lies in the binder development and in the optimization of the properties of the feed-stock material. Binder development has traditionally been a complex issue since it is related not only to the binder formulation but to optimization of the removal conditions as well. Therefore, mechanistic understanding of the physico-chemical processes that occur during thermolysis and information related to the rheology of the binder system are needed. In addition, the FDC method imposes several requirements for the binder system.In this study, the requirements of a binder system to be used in FDC are classified into two categories: general and secondary. The general requirements are low melting point, low viscosity and tackiness.

The general requirements are low melting point, low viscosity and tackiness. The secondary requirements are those that gain importance during specific processing stages (e.g.; extrusion, debinding, etc.) and are wetting ability (YSL and Ysv), melt strength, melt index, shear-thinning behavior, handling strength, (steric) stabilization ability, residue content, etc.A typical processing sequence for FDC involves milling of the ceramic powder, mixing the organic additives with the powder, extrusion of the feed-stock filament, solid freeform fabrication, binder removal and.sintering. For example, shear-thinning behavior and appropriate melt strength are required during the extrusion of the small-diameter « 2 mm) filament. The ~.J:).~. is the only ceramic shaping method that requires no die or mold,· and,>hence, the liquid-air interfacial energy·of the binder system determines the surface finish of the sintered body.

The ~.J:).~. is the only ceramic shaping method that requires no die or mold,· and,>hence, the liquid-air interfacial energy·of the binder system determines the surface finish of the sintered body. The requirement for the level of the paste viscosity upon melting.is perhaps the lowest one among the ceramic shaping techniques. For example, in injection molding and extrusion, the acceptable661viscosity is higher thanthaffor the FOC,sincethe former methods em.ploy force for shaping the component. However, low viscosity of the bipder above the melting point can cause deformation ("slumping") during the binder removal stage [2]. Below the melting point, strength of the binder is importi:\pt.forhandling thefilam~nt..Thesereq\liretnents litnit the binder formulation to a thermoplastic one. Thermoplastic polymers can besemicrystalline(e.g.; PolY9t~ylepe'Bol~mpylene,nylon) or glassy (e.g.; polystyrene, polyvinyl chloride).

Thermoplastic polymers can besemicrystalline(e.g.; PolY9t~ylepe'Bol~mpylene,nylon) or glassy (e.g.; polystyrene, polyvinyl chloride). Asernicrystalline polymer becomes ~iscous liquid when.tnelted, whereas a glassy polymerbecol11es a leathery or viscousdOugh-likematerial.when heated above its glass temperatur9' Tg [3,4]. Thus, semicrystalline thermoplastic polymers are suitable as organic additives in FOC.Since lowviscosityis desired upon melting, wax has been the major compo~ent of the binder system. Wax is a general~ame for a variety ofc~mpoundsthat include paraffinic waxes, microcrystalline waxes, synthetic hydrocarbon Waxes and oxidized polyethylene waxes [5]. Paraffinic and microcrystallinewaxt'tsarebr-products of the petroleum industry, whereas synthetic hydrocarbon.waxes and oxidized polYt'tthylene waxes are :produced by polymerization or oxidation of ethylene, respectively [5].

I3Y adjusting the parameters during the polymerization of ethylene, not only the chain length (as represented by average molecular weight and molecular weightdi~tribution), but also the degree ofbranching in the polymer (which determines the crystallinityand density) can be adjusted.The very low molecular weight (as low as 400) polyethylene•.(PE) is used as synthetic wax, whereas the very high end of the range (as high as 5 million) is known as ultra hi~hMwPE.. The degree of branching, on the. other hand,. determines the crystallinity, and thus, the density of the polymer.The low density polyethylene (LDPE) is highly branched and has about 50 % crystallinity. The high density form (HDPE) has few side chain branches and has about 90 % crystallinity[6, 7].Olefins(propylene, I-butene, I-hexene, l-octene) are copolymerized to improve selected properties of HDPE as the density is reduced.

Low density copolymers containing varying amounts of vinyl acetate (EVA), maleic anhydride (EMA), ethyl acrylate (EEA) are produced to modify selected physical properties of LDPE homopolymers [6,7].Properties of wax can be improved by the addition ofPE homop(jl~meror copolymers containing PE. This blending process not only improves the propertie~ofthewax, but also yields a graded volatility binder system. Despite the factthat every polymer has some solubilityin every other polymer, the magnitude in most cases is extremely low [8, 9]. Compatibility of the wax with EVA depends on the melting points of the two, and on the vinyl acetate content of the copolymer.In ceramic processing, organic additives notonlyaid in the shaping of the coll1Ponent, but also have significant influence on the microstructure, and thus, the properties of the ceramic component.

The purpose of this study is to apply general ceramic processing science to FOC, by evaluating different grades of EVA-petroleum wax mixtures as a potential binder. The fact that EVA and the petroleum wax are commercially available in several different grades, make them662good candidates to be used as binder components for FDC; therefore, a discussion on the effect of their properties is given in this study.II. ExpeI"im.entaIProcedureViscosityofthe EVA~waxmixtureswas measured by using a Brookfield viscometer with spindle #2. Since thermal gradients develop during.non--isotherml:jlviscosity measurem,entin:viscqus fluids, aSample holder which is slightlylarger thanthe/BroOkfieldtY2spindle indiameter,was used. A coolinglheating ra.teof30°C/hour was used, while the rotational speed of the spindle was selected depending on the viscosity range of the blend.

A coolinglheating ra.teof30°C/hour was used, while the rotational speed of the spindle was selected depending on the viscosity range of the blend. Penetration tests were performed on the I g samplespf the blends by using a 0.8g/mm2 contactpressure ala heating rate of 15°Clhour in a specially made penetratioiltestapparatus. Melting endothennsofdifferent grades ofthe EVA, wax,and their blends were characterized by using a Perkin-Elmer:DSC~4differential scanning calorimeter at 10°C/rrrln heatingrate,.with 5lng samples, and under flowing Nz. The temperature scale was calibrated by using the melting endotherm of indium sample. Microstructural analysis of the extruded filaments wetedone by using a Hitachi S~800 SEM equipped with a thermoionic filament, at 10kV accelerating voltage.The EVA used was supplied by DuPont andwas a copolymer of vinyl acetate and polyethylene, whereas the microcrystalline wax used was supplied by IGl.

The selected properties of the materials used as claimed by the supplier are given in table 1. The ceramic powder used was Alcoa A~17 grade alumina. The ceramic green bodi.eswere fabricated by using a Stratasys FDM 3~DModeler.Table 1: Selected properties of the ethylene~vinylacetatecopolymer grades used, as claimed by the /supplier.V.A. wt. % 20 20 20 25 28 28 28 28 28 28 28 28 28 28M.l. (dg/min) 500 8 2.5 400 2500 800 400 150 43 25 6 3 0.5 0.15Avg. Mw49100132000 51400 29990 40840 4401099360 112000 81900 119900Tensile Strength (MPA) 4.7 18 23 3.3 1.6 2.6 2.8 5.5 9.7 11 24 29 22 23663III. Results and·DiscussionsIn formulation of a binder system for FOG by using wax and EVA, the main factor that determiJJ.esthehandlings~rength0fthefilament feedstockare the melting points of the two components.

To obtain good strength, the melting point of the weaker component (wax) needs to be higher than that ofthestrongerone·.(EVA).During cooling, th~wax solidifies first and forms microglobe~ in the EVA melt,and eventually the polymer solidifies and forms a 3-dimensional network. During the early experiments inthis study, it was observed that paraffinic wax-EVA blends don't display handling strength due to the low melting pointof the paraffin wax. On the other hand,the melting point of the EVA can be adjusted by the vinyl acetate content, as shown in figure 1. The vinyl acetate disrupts crystallinity of the ethylene chain andde.creases the melting temperature of the polymer.Atagiven vinyl acetate content, the rheological properties of EVA can be adjusted by changing the molecular weight of the polymer.

As showninfig?re 3, the EVA containing 20 % vinyl acetate showed compatibility as the weight of the EVAwas varied from 5 to 30 %. The fact that the slopes of the curves in the 5-20 % range are almost identicaJ, reveal that the EVA can form a three dimensional network in the solvent wax and can dominate the viscoelastic/viscous flow. The increase in the activation energy at 30 % can be due to the entropic contribution.Resistance to viscous flow is especially important during the binder removal step, since binders with too low viscosity can cause slumping of the green body. During the thermolysis, since the wax is removed at an earlier stage, the EVA content in the remaining of the binder will increase,664and viscosity decrease of the binder system due to temperature increase will be compensated to some extent.

Figure 4 shows that the lower temperatures the wax is removed, the higher the0.8 - , - - - - - - - - - - - - - - - - - - ,0.70.60- ~ 0.5 .s 0.4 ~ it 0.3 Cii ~ 0.20.12500 - - 800 --·400 --_. 150 ........... 25 - - 6 _.- 30.030405080 70 Temperature (oC)6090100110Figure 2: The effect of molecular weight(as represented by the melt index) on the melting endotherm of EVA, containing 28 % vinyl acetate.30%20%15%10%5%0.0060.0070.0080.0090.0100.0110.0121 IT (oC)Figure3: The Arrheniusplot ofthe viscosity change ofthe microcrystalline wax - 20 % vinyl acetate containing EVAblends during heating and cooling at a rate of I.SoC/min. The increase.in the hysterisis is due to the increase in viscosity, which in tum causes high thermal gradients in the sample cup during heating or cooling.665increase./in·the visc.osity.ofthebinder system.

The viscosity••·of the billder system, together with the pore size distribution, will alsoaffectt~~r~distributionofthebinderwithin the green body during the binder removal. It was stated that binder will be removed from larger pores first and it will bere.distributed towards the sma!lerpor¢s due to their high suction potential [2].The diUlenSionalchange ofthebinder is also an important issue, since it can cause stress build up orwarpingin the green body.The thermal exp~nsion ofmicroscrystalline wax-EVA (20 % vinyl acetate) blends, which continues till the onset of the melting, is shown in figureS in the form of a penetration test. In figureS, the fact that the penetration ~ecomesmoregradual as the EVA content incre;;tses reveal that m~lt strength increases as the EVA content is increased.

In figureS, the fact that the penetration ~ecomesmoregradual as the EVA content incre;;tses reveal that m~lt strength increases as the EVA content is increased. This is to be expected, and a gradual decrease in the melt strength is desirable during the extrusion of the filament feedstock.16000 - , - - - - - - - - - - - - - - ,14000at 100°C12000£' ~ 10000..at 140°C 4th power regression'&:-808000,~ -g 6000 :2 ~ 4000 OJ2000oo5101520253035Weight % EVAFigure 4: The effect of the EVA content on the viscosity of microcrystalline wax - 20 % vinyl acetate containing EVA blend.A microstructural evaluation on theJractl1re surfaces oftliefilamentfeedstockscontaining60 % alumina support the previous arguments. The samples contained 24 % microcrystalline wax and 16 % EVA which had 24 %vjnyl acetate on average.

The samples contained 24 % microcrystalline wax and 16 % EVA which had 24 %vjnyl acetate on average. As shownin figure6,A,the binder that contains the higher molecular weightEV,A forms a bulkier network, thus providing better handling strength as compared to the binder with lower molecularweightEVA. Since all organic additives usedin ceramic processing are t~mporaryadditives,their removal characteristics from the green body is as important as their functional properties. The pyrolysis characteristics of the binder systems in EVA· wax system developed for FOC is discussed in else where [10].IV. ConclusionsBinder formulation is a complex issue and several factors need to be considered for intelligent formulation.

ConclusionsBinder formulation is a complex issue and several factors need to be considered for intelligent formulation. Blends in the wax-EVA system have been evaluated as potential binders to be usedin fused deposition of cer~cs.The factthat EVA and wax are available in several different grades make this system an excellent candidate as a potential bincl,er forBDC, and the effect of some.of the parameters have beendiscussedin thiS study. In order to obtain good handling, . -__. , - -_ . , - -_ . , - ----r 100Temp.906 eL- l!! ::l iii (jj a. E <ll....800.20.0303540Time (min.)Figure 5: The effect of EVA content on the penetration~estresult of a microcrystalline wax - 20 % vinyl acetate containing EVAFigure 6:Fracturesurfacesof(A)alurt1ina~reenbocl,y (B) alumina green body thatcontains·high·molecular welQ:nt667strength, the melting point of the polymer needs to be lower than that of the wax.

In this context, it was shown that as the vinyl acetate contentin theEVA increases, the melting point of the polymer decreases. By measuring the viscosity as a macroscopic property, it was shown that 20 % vinyl acetate containingE"Aisllliscible in microcrystalline waxupto,at least, 30 %. Binders used inFDC needs to have low· viscosity but their use may cause slumping during binder removal. Thus., thermolysis ofwax at low temperature is suggested as a solution. The fact that vohun.etric change upon heating increasesa& the EVA content in the binder increases was displayed in the form of a penetration test, which also shows the increase in the melt strength (or softening point) as the EVA content in the binder increases.Finally, the formulation of a binder system depends not only on the room temperature characteristics, but also on the degradation behavior of the binder.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperBINDER JET ADDITIVE MANUFACTURING OF STAINLESS STEEL - TRICALCIUM PHOSPHATE BIOCOMPOSITE FOR BONE SCAFFOLD AND IMPLANT APPLICATIONSKuldeep Agarwal, Sairam Vangapally and Alexander SheldonDepartment of Automotive and Manufacturing Engineering Technology, Minnesota State University, Mankato, 56001AbstractScaffolds are 3D biocompatible structures that mimic the extracellular matrix properties (mechanical support, cellular activity and protein production) of bones and provide place for cell attachment and bone tissue formation. Their performance depends on chemistry, pore size, pore volume, and mechanical strength. Recently, additive manufacturing (AM) has been used as a means to produce these scaffolds. This paper explores a new biocomposite manufactured using Binder Jet AM process.

Recently, additive manufacturing (AM) has been used as a means to produce these scaffolds. This paper explores a new biocomposite manufactured using Binder Jet AM process. Stainless steel and tricalcium phosphate are combined to form a composite and used in different volume fractions to produce parts with varying densities. Layer thickness, sintering time and sintering temperature are varied to study the effect of process parameters on the microstructure, dimensions and mechanical properties of the resulting structure.

It is found that the resulting biocomposite can be tailored by varying the process to change its properties and mimic the properties of scaffolds in bone tissue applications.IntroductionBiomedical implants can be classified into 3 major categories: external to the body (non- clinical, which includes surgical instruments, prosthetics etc.), internal to the body & permanent (includes hip implants, knee implants, stents etc.), and internal to the body & temporary (includes scaffolds, degradable screws and drug delivery systems) [1-3, 14].According to FDA, a ‘‘permanently implantable device is a device that is intended to be placed into a surgically or naturally formed cavity of the human body for more than one year to continuously assist, restore, or replace the function of an organ system or structure of the human body throughout the useful life of the device.’’ Some examples include knee and hip implants.

Temporary implants are commonly used in sports surgeries, such as in shoulder and knee ligamentous reconstruction and spinal reconstructive surgery [5, 6, 14].Scaffolds are temporary porous structures implanted to assist in tissue or bone regeneration. They are three-dimensional structures that mimic the extracellular matrix (ECM) properties (mechanical support, cellular activity, and protein production) and provide place for cell attachment and bone tissue formation. "The performance of scaffolds" depends on chemistry, pore size, pore volume and mechanical strength. Interconnected porosity is important for continuous ingrowth of bone tissue. Open and interconnected pores allow nutrients and molecules to transport to inner parts of the scaffold.

Interconnected porosity is important for continuous ingrowth of bone tissue. Open and interconnected pores allow nutrients and molecules to transport to inner parts of the scaffold. An ideal scaffold must satisfy the following requirements:23761Biocompatibility  Biodegradability  Appropriate porosity, pore size and pore shape  Bioactive  Mechanical strength  Adequate surface finish  Easily manufactured and sterilizedThe most common biomaterials for implants are metals and alloys, ceramics, and polymers. Among the metals, one of the most important classes of materials due to its corrosion resistance is stainless steel. Bones have elastic moduli 7 - 30 GPa, yield stress 30 - 70 MPa, compressive strength 100 - 230 MPa, and tensile strength 70 - 150 MPa. The first generation of implants focused on replacement of the bone with a metal implant.

The first generation of implants focused on replacement of the bone with a metal implant. However, mechanical properties of metals differ considerably from natural bone: Stainless steel elastic modulus = 190 GPa, yield strength = 220 - 1213 MPa, tensile strength 586 - 1352 MPa [3, 4, 7].These differences lead to stress shielding resulting in loosening of the implant due to degradation of human tissues around implants and, consequently, further surgeries to replace the implants. Ceramics are inorganic materials with high compressive strength and biological inertness that make them suitable for scaffolds used in strengthening or replacing damaged bones and tissues. The most commonly used bioceramics are metallic oxides (e.g., Al2O3, MgO), calcium phosphate (e.g., hydroxyapatite (HA), tricalcium phosphate (TCP), and octacalcium phosphate (OCP)), and glass ceramics (e.g.

Bioglass, Ceravital).Calcium phosphates have the best biocompatibility and properties closest to natural bones: elastic modulus 7 -13 GPa, compressive strength 350 - 450 MPa, tensile strength 38 - 48 MPa and flexural strength 100 - 120 MPa. However, they have poor fracture toughness and tensile strength that limits their application to bioimplants. Several in vitro and vivo works have shown that calcium phosphates support the adhesion, differentiation, and proliferation of osseogenesis-related cells (e.g., osteoblasts, mesenchymal stem cells), besides inducing gene expression in bone cells. The most important calcium phosphate is hydroxyapatite (HA, Ca10(PO4)6(OH)2) with chemical characteristics similar to hard tissues such as bone and teeth, that promotes hard tissue ingrowth and osseointegration when implanted into the human body. The porous structure of this material can be tailored to suit the interfacial surfaces of the implant.

The porous structure of this material can be tailored to suit the interfacial surfaces of the implant. As a bulk material, HA lacks sufficient tensile strength and is too brittle to be used in most load bearing applications. In such cases, HA is coated onto a metal core or incorporated into polymers as composites. The ceramic coating on the titanium implants improves the surface bioactivity but often fails as a result of poor ceramic/metal interface bonding [8-11].𝛂-TCP and β-TCP are the two crystalline varieties of HA of interest in biological applications. β-TCP is the thermodynamically stable form at low temperature. It transforms into 𝛂-TCP in the temperature range 1120 - 1170 oC. β-TCP is generally preferred in sintered ceramic implants, while 𝛂-TCP is more commonly used in bone graft cements because of its hydrolysis properties.

β-TCP is generally preferred in sintered ceramic implants, while 𝛂-TCP is more commonly used in bone graft cements because of its hydrolysis properties. The requirements that allow bone ingrowth are a porosity of 30 - 70 vol % and a pore diameter between 300 and 800 µm, mechanical properties of 0.5 - 15 MPa similar to cancellous bone.23772Thus there is a drive in the biomedical industry to create novel materials that behave very similar to bone and can be used for multiple applications: from permanent to temporary implants. Furthermore, these materials need to be manufactured in a manner that would create porosity in situ for biological applications. The current techniques used for making scaffolds are salt leaching, gas forming, phase separation, freeze-drying.Due to the versatility of additive manufacturing, it is gaining a lot of popularity in the field of bone implants.

Selective Laser Sintering, Selective Laser Melting, Electron Beam Melting and Binder jet manufacturing have all been used to create various porous structures for biomedical implants [7, 9]. To accomplish the various different requirements of the implants it is necessary to create biocomposites that can have the strength properties of metals as well as the biological properties of bioceramics.The present work aims at creating such structures by manufacturing and studying a stainless steel (SS) – Tricalcium phosphate (TCP) biocomposite with binder jet AM process. In this work SS and TCP are combined to form a composite and used in different volume fractions to produce parts with varying densities.

Layer Thickness, Sintering time and Sintering temperature are varied to study the effect of process parameters on the microstructure, dimensions and mechanical properties of the resulting structure.Binder Jet-based Additive ManufacturingThe main technique of manufacturing using the binder jet process is as follows: (a) the CAD file is sliced into layers and a STL file is generated, (b) each layer begins with a thin distribution of powder spread over the surface of a powder bed, (c) using a technology similar to ink-jet printing, a binder material selectively joins particles where the object is to be formed, (d) a piston that supports the powder bed and the part-in-progress lowers so that the next powder layer can be spread and selectively joined, (e) this layer-by-layer process repeats until the part is completed, (f) following a heat treatment, unbound powder is removed and the metal powder is sintered together. Fig. 1 shows the details of the whole process.Fig.

Fig. 1 shows the details of the whole process.Fig. 1: Schematic of the Binder Jet Process (Courtesy The ExOne Company)23783Process parametersThe binder jet process described above can be divided into 3 basic steps: 1) binding, 2) curing and 3) sintering. There are various process parameters that can be changed to obtain a customized part in each of these steps. These include powder size, layer thickness during binding, part orientation in bed, heater power, roller speed, curing temperature, curing time, sintering time, sintering temperature, and sintering atmosphere. To study the effect of each of these parameters and their interactions would require a huge experimental design and many hundreds of samples and testing.

To study the effect of each of these parameters and their interactions would require a huge experimental design and many hundreds of samples and testing. A feasibility study was done to see if a biocomposite can be created using this manufacturing technique where only the volume fraction of TCP, layer thickness, sintering time, and temperature were varied.Experimental PlanThe materials used in the study were stainless steel 316 (SS316), which has a mean particle size of 30μm and an apparent density of 2.75 g/cc. The chemical composition of SS316 is shown in Table 1.C 0.08 max 2.0 maxMnS P 0.03 max 0.75 0.045 max max Table 1: Chemical composition of SS316 (wt %)SiCr 16.0 - 18.0Ni 10.0 – 14.0Mo 2.0 – 3.0β-TCP was obtained from Sigma-Aldrich (21218) with a mean particle size of 5μm and an apparent density of 1.92 g/cc. Scanning Electron Microscope (JEOL JSM-6510MV) was used to look at the sizes and distribution of the powders before the process.

Scanning Electron Microscope (JEOL JSM-6510MV) was used to look at the sizes and distribution of the powders before the process. The images are shown in Fig. 6.SS316 was used as a benchmark and is called Sample Set 1. The two powders were mixed in 2 different volume fractions: Sample set 2 with 80% SS316-20% TCP and sample set 3 with 60% SS316-40% TCP. The mixed powders and the SEM images are shown in Fig. 2 and 3, respectively.Fig. 2: SEM Micrographs for SS316 and TCP23794The 3 sample sets were then used as the input powders in the binder jet additive manufacturing. To understand the effect of process parameters of the binder jet process on the material properties, 2 different layer thicknesses were chosen: 50 μm and 100 μm. The sintering time was varied as 2 hours and 4 hours and the sintering temperature as 1100 oC and 1200 oC. The design of experimental matrix is shown in Table 2. Three replicates of each experiment were done.Fig.

The design of experimental matrix is shown in Table 2. Three replicates of each experiment were done.Fig. 3: Mixed sample set 1 of SS316-TCP (left) and the SEM image (right)Expt No. 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24% TCP Layer Thickness0 0 0 0 0 0 0 0 20 20 20 20 20 20 20 20 40 40 40 40 40 40 40 40(μm) 50 50 50 50 100 100 100 100 50 50 50 50 100 100 100 100 50 50 50 50 100 100 100 100Sintering Time (Hrs) 2 2 4 4 2 2 4 4 2 2 4 4 2 2 4 4 2 2 4 4 2 2 4 4Sintering Temp (0C) 1100 1200 1100 1200 1100 1200 1100 1200 1100 1200 1100 1200 1100 1200 1100 1200 1100 1200 1100 1200 1100 1200 1100 1200Table 2: Design of Experimental Matrix23805Printing, Curing and Sintering.Samples were printed based on ASTM E9 specification for compression testing. Three samples from each set were printed for compression testing. The roller speeds were kept at the minimum of 1 mm/sec.

Three samples from each set were printed for compression testing. The roller speeds were kept at the minimum of 1 mm/sec. The samples were cylinders with diameter of 12.5 mm and height of 22.5 mm.After printing the samples were cured in an oven at 1750C for 3 hours. The cured parts were sintered in an Ar atmosphere. After sintering the samples were removed and cleaned. The samples at each stage are shown in Fig. 4.Fig. 4: Samples while being printed and after the sintering and after machiningExperiments: Compression testingThe samples were tested according to ASTM E9 standards [10]. The samples were setup for compression testing on the MTS 810 material testing system. The MTS machine was set to run failure detect mode test at speeds of 15 mm/min. The stress strain curves were then calculated for all the experiments. The sample in the machine for testing and the crushed sample after testing are shown in Fig. 5.Fig.

The stress strain curves were then calculated for all the experiments. The sample in the machine for testing and the crushed sample after testing are shown in Fig. 5.Fig. 5: Compression testing23816Microstructural AnalysisScanning electron microscope was used to study the microstructure of samples before and aftersintering. The images were used to calculate the neck to diameter ratio for each sampleNeck size ratio =𝑥𝐷Where, x is neck diameter and D is particle diameter. Ten different measurements were takenfrom different images of the sample and they were averaged. The images were also used to calculate the pore sizes. The behavior of TCP and its interaction with SS316 was also studied [12, 13, 14].Results SEM AnalysisThe sintered SS316-TCP biocomposites were examined under a SEM. The resultingstructures are shown in Fig. 11 with both the constituents.Fig.

The resultingstructures are shown in Fig. 11 with both the constituents.Fig. 6: SEM Images of (a) Expt 1 (b) Expt 2 (c) Expt 3 (d) Expt 4 (SS316 with 0% TCP at 50 μm Layer Thickness)23827TCPSS316TCPFig. 7: SEM Images of the SS316-TCP biocomposite. (Left) – 80% SS316-20% TCP by volume, (Right) – 60% SS316-40%TCP by volume. Both are for 50 μm layer thickness, 2 Hour sintering at 1100 0C)Fig. 8: Variation of densities in different experiments (LT = Layer Thickness)Figure 6 and 7 show the different stages of sintering and the effect of different compositions. The experiment 1 (time = 2 hr and temp =1100 0C) has a very small neck to diameter ratio showing the onset of the sintering (initial phase). Comparison between experiments 1 and 2 and experiments 1 and 3 shows that temperature has a greater effect on sintering than the time. The pore size also decreases from an average of 20 μm in experiment 1 to 8 μm in experiment 4.

The blank spaces indicates that the samples crumbled immediately on compression238611The compression tests show the following features: a) 50 Micron Layer thicknesses have higher compressive strength than the 100-micron layerthickness for all material compositions.b) The strength follows the same pattern as the densities and the neck – diameter ratios.

Thestrength is higher for 4 hour 1100 0C than for 2 Hour 1200 0C sintering.c) The samples were not sintered completely at 2 hours and 1100 0C and hence crumbledimmediately on compression.d) The compressive strength varied from a peak of 1987.7 MPa for certain settings to a low of803.8 MPa for SS samplese) The compressive strength of 20% TCP samples was higher for the same settings than the40% TCP samples.f) The Elastic modulus of both 20% and 40% TCP samples is within the range of cortical andcancellous bones (7-30 GPa).Thus, there are a variety of bone implant and scaffold applications that are possible by a biocomposite manufactured by binder jet AM process. For some implant applications 20% TCP- 80% SS biocomposite can be used, while for others a 40% TCP – 60% SS biocomposite can be used.

For some implant applications 20% TCP- 80% SS biocomposite can be used, while for others a 40% TCP – 60% SS biocomposite can be used. The processing can be done differently to make sure that the materials have the required strength.ConclusionsScaffolds are 3D biocompatible structures that mimic the extracellular matrix properties (mechanical support, cellular activity and protein production) of bones and provide place for cell attachment and bone tissue formation. Their performance depends on chemistry, pore size, pore volume, and mechanical strength. Recently, additive manufacturing (AM) has been used as a means to produce these scaffolds. This paper explores a new biocomposite manufactured using Binder Jet AM process. SS and TCP are combined to form a composite and used in different volume fractions to produce parts with varying densities.

SS and TCP are combined to form a composite and used in different volume fractions to produce parts with varying densities. Layer thickness, sintering time, and sintering temperature are varied to study the effect of process parameters on the microstructure, dimensions, and mechanical properties of the resulting structure. It was found that the resulting biocomposite can be tailored by varying the process to change its properties and mimic the properties of scaffolds in bone tissue applications.The different compositions can be used to produce implants or scaffolds depending on their strength characteristics. 20% TCP biocomposite is more suitable for implant applications, while the 40% TCP biocomposite is more suited for scaffold applications.Further work needs to be done to determine the biocompatibility and wear characteristicsof these materials before can be a candidate for the biomedical applications.References1. Tarafder, S., 2013.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperBINDER JETTING ADDITIVE MANUFACTURING OF WATER-ATOMIZED IRONI. Rishmawi*, M. Salarian*, and M. Vlasea* *Department of Mechanical and Mechatronics Engineering, University of Waterloo, ON, Canada N2L 3G1AbstractBinder jetting additive manufacturing (BJAM) was deployed to processing of low-cost, water-atomized pure iron powder. Surface morphology and particle size distribution of the powder were fully characterized using scanning electron microscopy (SEM) and particle dynamic image analysis via Retsch Camsizer X2. Cylindrical samples were fabricated, and in the AM process, the effects of powder compaction, layer thickness and liquid binder level on green part density were studied. Density analysis was performed using x-ray computed tomography (µCT).

Density analysis was performed using x-ray computed tomography (µCT). The potential application and future research work will be outlined based on the characterization results.Key words: Binder jetting, Additive manufacturing, Irregular iron powder, Sintering schedule, Part densityIntroductionBinder Jetting Additive Manufacturing (BJAM), also known as 3D Printing (3DP), is a powder bed additive manufacturing (AM) process in which powder is spread in layers across a build bed and liquid binding ink is jetted onto each layer. BJAM is a two-step process that goes from initial materials to final part. In the first step, geometrical features of the part are achieved using AM. In the second step, functional features of the parts are achieved via sintering. The process has been predominantly used with ceramics, composites and metals, with metals receiving increasing attention.

The process has been predominantly used with ceramics, composites and metals, with metals receiving increasing attention. BJAM of metals has been utilized with iron- based alloys [1–13] for several target applications, such as biomedical implants [3], [6], forging dies [7] and injection molding tools [8,9].Aside from advantages such as customizability and absence of thermal stresses in printed specimens, BJAM is also a comparatively low-cost process. This is mainly due to the ability to print at room temperature conditions and batch sinter the printed specimens. There has been recent interest in water-atomized powders for BJAM process, which can produce specimens of properties comparable to gas-atomized powders [14,15]. Mostafaei et al. [14] studied the densities and microstructures obtained from water-atomized Inconel 625 specimens and report densities of up to 95% after sintering. Inaekyan et al.

Mostafaei et al. [14] studied the densities and microstructures obtained from water-atomized Inconel 625 specimens and report densities of up to 95% after sintering. Inaekyan et al. [5] characterized density and mechanical properties of water-atomized pure iron specimens, but densities of only around 45% are reported after unassisted sintering. However, the authors performed sintering at 1100°C, which is significantly lower than the melting point of the material [16].This study focuses on printing BJAM samples with water-atomized iron powder. Apart from feedstock powder characteristics, the most significant process parameters are powder compaction and spreading, layer thickness and liquid binder level. These three parameters are studied to maximize green part density. Subsequently, the impact of sintering on final part density and surface quality is studied.

These three parameters are studied to maximize green part density. Subsequently, the impact of sintering on final part density and surface quality is studied. The importance of this work is in bridging the gap between BJAM and large-scale industrial adoption, where inexpensively prepared feedstock powders produce high-quality parts.ExperimentalMaterials The powder used is water-atomized pure iron (AT-1001-HP, Rio Tinto QMP). The powder is sieved to a size range of < 45 μm (USA Standard Test Sieve No. 325, Hogentogler & Co.). Size distribution is characterized (Retsch Camsizer X2) and surface morphology (LEO Field Emission SEM 1530). The powder is blended with polyvinyl alcohol (PVA) powder (Alfa Aesar) that is previously ground and sieved to size range of < 63 μm. The160final feedstock blend is 99% iron and 1% PVA by mass.

The powder is blended with polyvinyl alcohol (PVA) powder (Alfa Aesar) that is previously ground and sieved to size range of < 63 μm. The160final feedstock blend is 99% iron and 1% PVA by mass. A water-based solvent (ZbTM 60, Z Corporation) is used as a liquid binder in the BJAM process.AM Process Parameters Printing is accomplished on a ZCorporation 310 Plus binder jetting system (3D Systems). Three AM process parameters are varied: powder compaction, layer thickness and liquid binder level. Powder compaction is varied by the activation or deactivation of the roller that spreads the powder from the feed bed to the build bed. Layer thickness and liquid binder level are varied through the printer settings. The parameters utilized are summarized in Table 1. Cylindrical specimens measuring 5 mm in diameter and 8 mm in height are printed. Each print batch consists of 16 replicates.

The parameters utilized are summarized in Table 1. Cylindrical specimens measuring 5 mm in diameter and 8 mm in height are printed. Each print batch consists of 16 replicates. The liquid binder level settings are presented in terms of pL/drop, where in terms of the ZPrint software, 70 pL/drop corresponds to 200% saturation level, and 105 pL/drop to 300%. The saturation level physically represents the amount of binder jetted.Table 1 Summary of process parameters and settings used for the experimental studyProcess ParameterLevelsParameter SettingsRoller actuation Layer thickness (μm) Liquid binder level (pL/drop)2 3 2On, Off 75, 100, 125 70, 105Post-process Parameters The heat treatment schedule is developed based on previous studies [5,14,17–19]. Sintering is performed on the specimen with the highest green density in order to obtain the highest possible sintered density [20]. An illustration of the schedule is presented in Figure 1.

Sintering is performed on the specimen with the highest green density in order to obtain the highest possible sintered density [20]. An illustration of the schedule is presented in Figure 1. The first stage of the heat treatment is de-binding, where the binder contained in the specimen burns out. The temperature is increased at a rate of 5°C/min and de-binding is conducted at 300°C for a duration of 30 min. In the second stage, the temperature is increased at 10°C/min and held at1490°C for 6 hours.Figure 1 Heat treatment protocol applied to the specimensCharacterization Green and sintered densities are determined via x-ray computed tomography (μCT) (Zeiss Xradia 520 Versa) at a voxel size of 5.5 μm and a voltage of 140 kV. Image processing and porosity analysis are performed161(Dragonfly Pro v3.3, Object Research Systems Inc.).

Image processing and porosity analysis are performed161(Dragonfly Pro v3.3, Object Research Systems Inc.). A photographic representation of a specimen in the green and sintered states is shown in Figure 2.Figure 2 Specimens as a (a) CAD image, and in the (b) green and (c) sintered statesShrinkage Dimensional shrinkage in the specimens after sintering is determined through equation (is the measured dimensional feature of a green specimen and is that of the specimen after sintering. Geometrical measurements are performed by using digital calipers (Mitutoyo Absolute Digimatic Caliper). Three measurements are taken for each dimension of each specimen, and the measurements are performed at approximately the center of the specimen for both diameter and height.), where1𝑥𝑥𝑆𝑆()𝑥𝑥𝐺𝐺−𝑥𝑥𝑆𝑆𝑆𝑆ℎ𝑟𝑟𝑟𝑟𝑟𝑟𝑟𝑟𝑟𝑟𝑟𝑟𝑟𝑟 𝑟𝑟𝑟𝑟 𝑥𝑥 =𝑥𝑥𝐺𝐺 × 100Surface Roughness Surface roughness of green and sintered specimens is measured by using a laser confocal microscope (Keyence VK-X250).

Scans of the side and top surfaces of the specimens are performed. The measurements are taken from approximately the same locations on all specimens for consistency. Scans are performed at a vertical resolution of 1 μm, with a scan area of 1400 μm by 1000 μm.1Results and DiscussionPowder Characterization Powder size analysis results reveal a size distribution of 18 − 54 μm, with a D50 of 34.1 ± 0.5 μm. Powder sphericity at the D50 size is 0.8. SEM images of the powder are shown in Figure 3. The irregularity and angularity of the particles are visible in the figure.162𝑥𝑥𝐺𝐺Figure 3 SEM images of the iron powder at magnifications of (a) 200× and (b) 500×Green Density Analysis Specimen densities are determined through μCT scans and image processing. A consistent voxel size for both green and sintered specimens is used to ensure consistent analysis.

Based on the results in Table 2, powder compaction through roller actuation is clearly favorable for increasing green density. A powder bed that is more densely packed allows for higher density of the green specimen. The results show that a lower layer thickness leads to higher green densities when the roller is activated. A lower layer thickness corresponds to a tighter packing of the powder in the bed, as the roller spreads a thinner layer of particles an overall larger number of times. Furthermore, a lower liquid binder level leads to higher green densities with the roller activated. The liquid binder takes up space in between powder particles. Hence, minimizing its quantity should increase specimen density, provided that structural integrity of the specimen is maintained. Excessively high levels of liquid binder must be avoided as they might lead to layer shifting, in which individual layers of powder get shifted and results in a slanted specimen.

Excessively high levels of liquid binder must be avoided as they might lead to layer shifting, in which individual layers of powder get shifted and results in a slanted specimen. Roller actuation also affects powder163spreading. Observations during printing reveal that powder spreading is much less uniform when the roller is deactivated. Streaking and unevenness in spreading are noted, as well as powder clinging to the roller. Generally, layer shifting is occasionally observed when the roller is deactivated, but it is not observed at all when it is activated. Powder spreading and compaction and their optimization are interesting topics of research for future work.Sintered Density Analysis Sintering is performed on the specimen set producing the highest green density, namely, specimen A. A sintered density of 91.3% is obtained. Image processing analysis on the μCT scans reveals the distribution of porosity after sintering.

A sintered density of 91.3% is obtained. Image processing analysis on the μCT scans reveals the distribution of porosity after sintering. Three-dimensional renders of the specimen in the green and sintered states are illustrated in Figure 4.Figure 4 Three-dimensional renders of the samples in the (a) green and (b) sintered (1490°C) states.The final sintered density value of 91.3% shows that BJAM of water-atomized iron powder can produce high density parts while preserving shape fidelity for simple, symmetric shapes. The high temperature and duration used promotes aggressive sintering and densification by the formation of sinter necks among powder particles. Higher densities are expectedly attainable if longer sintering durations were used. Sinter neck formation reduces the total energy of the thermodynamic system, which is the driving force, and the reduction in energy is higher for a smaller powder particle size [20].

Sinter neck formation reduces the total energy of the thermodynamic system, which is the driving force, and the reduction in energy is higher for a smaller powder particle size [20]. Sintering progresses once an activation barrier is crossed, whereby volumetric diffusion is the major densification mechanism at such a high temperature [20]. The activation barrier for pure iron is between 1000°C and 1200°C [21]. Maintaining the sintering environment for longer durations at a constant temperature that is beyond the activation barrier promotes sinter neck formation. Thus, it is possible to select a sintering temperature that is between 1200°C and the melting point, and increase the sintering duration as long as needed in order to obtain the maximum desired part density. At such a temperature, sintering will progress and only cede when the total energy of the system is reduced to the minimum.The results presented in this study are comparable to those presented by Inaekyan et al. [5].

[5]. In their work, the authors deploy a powder material identical to that used in this work, but with a size distribution of 15 to 51 µm.164The sintering schedule resembles an industrial Powder Metallurgy schedule, whereby de-binding is performed at 195°C for 6 hours, and sintering at 1100°C for 2 hours in a pure argon atmosphere. The authors report green and sintered densities of ~40% and ~45% respectively. The final sintered density result is drastically different to what is obtained in our work. This difference underscores the importance of the sintering schedule. As discussed, the sintering to high densities occurs by crossing the activation barrier and maintaining the temperature for longer and longer durations. Sintering at a lower temperature would require longer durations to achieve high densities.

Sintering at a lower temperature would require longer durations to achieve high densities. The sintering atmosphere also plays a role, where the introduction of hydrogen as a reducing agent aids in preventing oxide layer formation, which would hinder sintering.Porosity distribution is presented as a color map in Figure 5, showing the 91.3% density sintered specimen from the side and top. Sintering to the density of 91.3% results in the absence of an interconnected pore network, where instead pores are isolated inside the specimen and not connected to the external air. Pore volumes in the range of 4.32×103 μm3 to 0.007 mm3 comprise the majority of porosity in the specimen. Larger pores are almost exclusively present in the topmost region of the specimen, while medium-sized pores are interspersed in the central region of specimen. Smaller pores are distributed throughout the whole specimen, with lower concentrations in the bottommost region.

Smaller pores are distributed throughout the whole specimen, with lower concentrations in the bottommost region. These observations are indicative of powder packing effects, where powder at the bottom of the bed is repeatedly compacted such that low porosity is obtained in the specimen bottom. Furthermore, pore distribution in the horizontal direction (perpendicular to the build direction) is uneven. Both small and large pores are located more preferentially toward one side than the other, which is an indication of a build orientation effect. Specifically, powder spreading might lead to preferential packing, where the one end of the build bed gets more powder than the other end and hence better packing.Figure 5 Three-dimensional and top-view flattened images of the cylindrical specimen sintered at 1490°C, obtained via μCT scans.

The distribution of pores by volume is shown as a color map.The work of this paper shows that high densities are attainable with water-atomized pure iron powder in a BJAM system. However, the necessary next step is the creation of a comprehensive and reliable “map” that allows165predictive control over sintering parameters to achieve pre-determined densities. This can be achieved through a Master Sinter Curve (MSC) [20]. The MSC would allow appropriate sintering schedules to be selected in order to obtain densities over 90% relevant to metal component production, but at lower temperatures than the temperature presented in this work (1490 °C) and at the cost of longer sintering durations.

Furnaces of lower temperature with environment gas control are economical and generally preferred by industry.Shrinkage and Surface Roughness Analysis Shrinkage induced on specimen A upon sintering is quantified by comparing the major and minor diameters of the cylinders a and b and height h before and after sintering. Shrinkage calculations are presented based on calipers measurements and illustrated in Figures 6 and 7. The shrinkage analysis shows that high shrinkage in all dimensions is observed. Shrinkage in a and b is 25.3 ± 2.8% and 24.2 ± 2.1% respectively, while in h it is 24.8 ± 3.5%. Furthermore, based on the CT data, shrinkage in a is 24.86 ± 1.37%, in b is 24.37 ± 2.42%, and in h it is 25.31 ± 2.37%.

Furthermore, based on the CT data, shrinkage in a is 24.86 ± 1.37%, in b is 24.37 ± 2.42%, and in h it is 25.31 ± 2.37%. Shrinkage is relatively uniform among the three dimensions, which indicates that shape fidelity is conserved for these simple, symmetric shapes despite the large reduction in volume.Figure 6 Visual illustration of shrinkage from the (a) green and (b) sintered samples166Figure 7 Shrinkage characteristics associated with the 91.3% density sintered specimenSurface roughness measurements are made on the green and sintered specimens, and the results are illustrated in Figure 8 against specimen density. The area-normalized arithmetical mean height, Sa, is used as a surface roughness metric. Correction for curvature is utilized as a post-processing step on the roughness values in the case of side surface roughness measurements. Side and top surface roughness values of 32.1 ± 3.4 µm and 21.4 ± 5.7 µm respectively are observed for the green specimen.

Side and top surface roughness values of 32.1 ± 3.4 µm and 21.4 ± 5.7 µm respectively are observed for the green specimen. By comparison, side and top roughness of 19.6 ± 3.4 µm and 18.0 ± 2.5 µm respectively are observed for the sintered specimen. Surface roughness drops by 16% − 39% upon sintering. Higher sintering temperatures promote densification, and the promoted formation of sinter necks results in smoother necking between particles and hence a lower roughness value. In general, top surface roughness is lower than that of the side surface, which suggests a build orientation effect.

In general, top surface roughness is lower than that of the side surface, which suggests a build orientation effect. Powder packing in the build direction might be looser than across the layers, due to compaction by the counter-rotating roller.167Figure 8 Surface roughness plotted against density for green and sintered specimensIn summary, this study shows that in-process parameters can be controlled in order to maximize green density, and sintering has an impact on porosity distribution, particularly in terms of interconnected porosity, and surface qualities. In-process and post-process parameters can be carefully selected to suit the application requirement and attain the required final density.ConclusionsWater-atomized pure iron powder was deployed in a commercial BJAM system and subsequently subjected to a sintering heat treatment in order examine the possible densities attainable. Green density can be increased by reducing layer thickness and liquid binder level.

Green density can be increased by reducing layer thickness and liquid binder level. Powder compaction has an important effect on powder spreading and packing in the bed, and optimization of powder spreading is a topic of future research. Sintering at a high temperature close to the melting point and for long durations can achieve densities up to 91.3%, and increasing sintering time is expected to produce even higher densities. In the context of industrial applications, high densities are expected to be attainable by sintering at lower temperatures but for extended durations, in industrial furnaces. The creation of a Master Sinter Curve is an important next step to allow predictive tailoring of densities. Shrinkage induced upon sintering is large, although it is uniform and shape fidelity is well-conserved.

Shrinkage induced upon sintering is large, although it is uniform and shape fidelity is well-conserved. Aggressive densification during sintering generally enhances surface quality of printed parts.AcknowledgementsThis work was supported by funding from the Federal Economic Development Agency for Sothern Ontario (FedDev Ontario), in partnership with Rio Tinto. The authors would like to acknowledge the contribution of Rio Tinto for their support and mentorship throughout the completion of work.168References[1] Y. Tang, Y. Zhou, T. Hoff, M. Garon, Y.F. Zhao, Elastic modulus of 316 stainless steel lattice structurefabricated via binder jetting process, Materials Science and Technology. 32 (2016) 648–656. doi:10.1179/1743284715Y.0000000084.[2] R. Frykholm, Y. Takeda, B.-G. Andersson, R. CarlströM, Solid State Sintered 3-D Printing Component by Using Inkjet (Binder) Method, Journal of the Japan Society of Powder and Powder Metallurgy. 63 (2016) 421–426.

nder Jetting of 316L process simulation tools evaluationK. Caballero1,2, V.A. Medrano 1,2, E. Arrieta1,2, H. Sandoval3, R.B. Wicker1,2, and F. Medina1,21 W.M. Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, TX 79968, USA  2 Department of Aerospace and Mechanical Engineering, The University of Texas at El Paso, TX 79968, USA 3 Lockheed Martin Missiles and Fire Control, Dallas, TX 75051, USAKeywords: Metal Additive Manufacturing, Binder Jetting, Sintering, Simufact Additive, ShrinkageABSTRACTBinder Jetting has become one of the most popular Additive Manufacturing technologies over the years due to its low cost and fast production times, nevertheless this technology has a steep learning curve due to the shrinkage induced to parts during sintering.

Since shrinkage is not uniform along the part, it’s hard to efficiently determine what areas will be distorted hence this needs to be taken into consideration when designing a new part and many iterations need to be printed until dimensional accuracy is achieved, as a result production time and cost significantly increase. New Binder Jetting simulation tools are being developed and tested; this software will help the technology be more robust and user-friendly for the industry. The software computes a sintering simulation and can provide displacement results making support positioning more efficient, in addition, newer versions of the software can export a compensated model which will be able to be sintered without supports. To evaluate the simulation software, a dimensional test artifact model was designed and printed, then compared with the software predicted model simulation results.

To evaluate the simulation software, a dimensional test artifact model was designed and printed, then compared with the software predicted model simulation results. The simulation software was used in an initial evaluation of the test artifact geometry to identify areas of concern in the model and document them so efficiency when predicting material behavior during the sintering process can be evaluated. In addition, an evaluation of the effects of different sintering process parameters on the physical and mechanical properties of the material will be analyzed considering the inert sintering atmosphere of the process.

Finally, printing parameters of the machine such as layer thickness, binder saturation, and recoat speed among others will also be evaluated.INTRODUCTIONMetal Additive Manufacturing is an important technology that enables the production of complex- shaped components and structures with high mechanical properties and a minimal material loss, which is a gradually improving the manufacturing sector [1].

Binder Jetting is an additive manufacturing technology that became popular in the early 1990s, it receives interest from the industrial communities because it allows the manufacturing of complex shaped components1519without involving the melting of the material and rapid solidification that can induce defects producing unwanted results [2].The Binder Jetting process consists of binder droplets being selectively deposited onto the powder bed, the powder with binder mix is then quickly cured and then another layer of metal powder is deposited on top of the previous layer, this process is repeated until the pre-liminary part called green part is completed. The green part then proceeds to be cured for 4 hours and then is carefully depowdered since it can easily break because the particles are not properly joined together yet.

The green part then proceeds to be cured for 4 hours and then is carefully depowdered since it can easily break because the particles are not properly joined together yet. To join the particles together, the green part will be subjected to a sintering process under an inert atmosphere to avoid oxidation, this process will properly bind together the metal particles.SIMULATIONSSimufact Additive is a powerful simulation software of metal additive manufacturing, recently, its manufacturer Hexagon added a new function to simulate the sintering process in Binder Jetting parts; distortion, shrinkage, residual stress, and temperature results can be provided by the software.

The software can also predict and compute a compensated model to counter the distortion on the printed part, producing a final product without the need of many iterations [4].Sintering ProfilePredefined time and temperature values were used, nevertheless material and green density of the printed part had to be manually inputed. Once the software has all the required information, the simulation will provide a result similar to the one that can be provided by an actual sintering process.MeshingThe mesh of the analyzed geometry is shown on the Figure 1, the overall dimensions of the artifact are 50 x 17 x 12 mm, the average element size is about 3 mm. This element size is based on the previous knowledge learned with simulations on the same software.Figure 1.

This element size is based on the previous knowledge learned with simulations on the same software.Figure 1. Finite element mesh of the artifact design1520MATERIALS AND METHODSPowderThe powder feedstock used in this printing work was gas-atomized 316L stainless steel supplied by Sandvik Osprey Powders with a spherical shape and a size range of 38 ± 5 microns, chemical composition is reported in Table 1.Fe 316L Balance ≤ 0.030CSi ≤ 1Mn ≤ 2P ≤ 0.045S ≤ 0.03Cr 16-18Ni 10-14Table 1. Chemical composition (wt%) of the used Stainless steel 316L powderThe powder goes through a process of optical characterization using a CAMSIZER X2 Microtac Retsch that provides precise particle size and shape information of powders in a measuring range from 0.8 μm to 8 mm, Figure 2 shows that Xc min displays the particle width, which is determined from the narrowest of all measured chords Xc, this characteristic is comparable with a sieve analysis [5].Fig 2.

Size definition of a particleTo define the shape of the particle the compactness must be calculated, this is determined by the particle length Xfe max and the particle area A. Generally, perfect circles or spheres have a compactness equal to 1, for all other shapes the compactness is < 1. The shape definition equation is calculated using Equation 1.Eq 1. Calculation of the Compactness1521Mo 2-3The results of the size analysis are shown in Table 2 and Graph 1, the stock powder had a particle size distribution of D10: 10.8 μm, D50: 19.3 μm, and D90: 32.1 μm.Table 2. Size definition of the powderGraph 1. Size distribution of the powderThe results of the shape characterization are shown in the Graph 2.Graph 2.

Size definition of the powderGraph 1. Size distribution of the powderThe results of the shape characterization are shown in the Graph 2.Graph 2. Shape characterization of the powder1522Finally, the powder was analyzed with a JSM-IT500 Scanning Electron Microscope [6] to observe particle shape under a microscope with a magnification of 50 μm, 20 μm and 5 μm in the Figures 3, 4 and 5, respectively.Figure 3. Microscope picture of 316L powder taken SEM Machine with a magnification of 50 μmFigure 4. Microscope picture of 316L powder taken SEM Machine with a magnification of 20 μmFigure 5. Microscope picture of 316L powder taken SEM Machine with a magnification of 5 μm1523Printing ProcessThe samples were printed using an Innovent+ metal 3D printing system manufactured by ExOne in 2018 as an upgrade version of the Innovent released in 2016, with a build volume of 160 x 65 x 65 mm [7].

The binder used was an aqueous-based solution provided by ExOne containing monobutylether, isopropanol and ethylene glycol. The first two evaporate during the curing while the ethylene glycol cross-links producing PEG [8].The printing parameters used for this build are reported in Table 3. These parameters have been selected based on previous printing experience.Drying Time (s)10Bed Temperature (°C)40Recoat Speed (mm/s)35Roller Rotation Speed (rpm)600Roughing Roller Rotation Speed (rpm) 300Roller Traverse Speed (mm/s)10Table 3. Printing Process ParametersAfter the model is fully printed, a curing process was carried out at 200 °C for 4 hours in air using a Yamato DX402C furnace.Post Process SinteringSintering bonds metal particles via mass transport mechanism and bring them to a denser state [9]. The loss of free surface energy during the sintering process can be explained by surface and bulk transport mechanisms.

Surface transport methods do not cause densification, bulk transport processes, on the other hand, cause neck expansion as well as shrinkage [10].For the sintering process we used a Carbolite Gero Model CTF12/TZF12 tube furnace in Argon atmosphere at 4 different temperatures, first, the temperature was raised at a rate of 5 °C per minute until 600 °C was reached, then it stays at that temperature for 30 minutes, the temperature is then increased at the same rate until it reaches 1000 °C and stays at that temperature for 2 hours, after this, temperature was increased to the final temperature of 1410 °C with a rate of 5 °C, and stays at that temperature for another 2 hours, finally, we decreased the temperature at a controlled rate, this slow decrement in temperature avoids drastic deformations on the part.Figure 6 shows the graph of the sintering profile, with a graphic demonstration of temperature against time.1524Figure 6.

Graph of the Sintering Process of Time against Temperature. Same one used for the Simufact simulationRESULTS AND DISCUSSIONSimulation ResultsSimulation results predicting shrinkage and deformation of the part are shown in Figure 7, and the real sintered part is shown in Figure 8, comparing these two figures, is clear that the simulated part has more deformation in the front face of the design, the hole and fin features also have greater deformation compared to the sintered part, the features got reduced by 12% of the initial dimension on the x direction. In addition, the part has a significant distortion in the thin walls, and some eventually broke due to this distortion.

In addition, the part has a significant distortion in the thin walls, and some eventually broke due to this distortion. The difference in our results can be due to several factors, such as the shape of the ceramic crucible on which the part was sitting during the sintering, this affects how the gas flow hits the green part and sintering orientation which varies the distribution of heat on the surface of the place with the surface of the green part.Figure 7. Simulation of the original part on the Simufact Software1525Figure 8. Sintered part before the compensationThe compensated part showed dimensional changes, shown in Figure 9. A dimensional change of 50 mm to 59 mm along the x direction was done to counter shrinkage and deformation of the part. Figure 10 shows the dimensional change and the quality in the fin features once the part has been subjected to the sintering process.Figure 9. Compensated part with the new dimensionsFigure 10.

V. Nagaraju, S. Kumaran, and T. Srinivasa Rao, “Microwave sintering of 316L stainless steel: Influence of sintering temperature and time,” Materials Today: Proceedings, vol. 27, pp. 2066–2071, Jan. 2020, doi: 10.1016/j.matpr.2019.09.062.ACKNOWLEDGEMENTSThe presented research was performed at The University of Texas at El Paso (UTEP) within the W.M. Keck Center for 3D Innovation (Keck Center). The authors are grateful to UTEP students Victor Medrano and Julio Diaz and UTEP professors Francisco Medina and Edel Arrieta for their contributions to the project. Primary support for this research was provided by Lockheed Martin Corporation as well as strategic investments via discretionary UTEP Keck Center funds and the Mr. and Ms. MacIntosh Murchison Chair I in Engineering Endowment at UTEP. The views and opinions expressed in this article are those of the authors and do not necessarily reflect the official opinion or policy of any agency of the US government (if applicable).1527.

Ferguson2, E. Shafirovich2, Y. Lin1 and Ryan B. Wicker11W. M. Keck Center for 3D Innovation. College of Engineering E-108. University of Texas at El Paso.University Ave. 500 W, El Paso TX. 79968, USA.2Department of Mechanical Engineering, The University of Texas at El Paso.Corresponding author:cdiazmoreno@utep.eduKeywords: Ceramics; Aluminum nitride; 3D components fabrication; Thermal conductivity;Binder jetting; Additive manufacturing;AbstractThis work reports on the novel fabrication of aluminum nitride (AlN) complex components usingbinder jetting, on the use of sintering and hot isostatic pressing (HIPing) to increase their density,and on the characterization of the printed material, including thermal conductivity. The HIPingparameters employed were a temperature of 1900 °C using a rich nitrogen atmosphere at apressure of 30,000 psi during 8 h.

The HIPingparameters employed were a temperature of 1900 °C using a rich nitrogen atmosphere at apressure of 30,000 psi during 8 h. Results show that the printed and HIPed AlN components hada 1.96 g/cm3 (60.12%) density when compared to theoretical values. The thermal conductivityfor densified and HIPed components was measured in the range from 23 °C to 500 °C resultingin values from 4.82 W/m*K to 3.17 W/m*K, respectively. Characterization using scanningelectron microscopy, energy dispersive X-ray spectroscopy and X-ray diffraction was used toinvestigate the ceramic structural morphology of the sintered and HIPed material, its chemicalcomposition, and crystal structure of the binder jetting manufactured AlN components.1. Introduction1.1 Introduction to Additive Manufacturing143Additive manufacturing (AM) is a novel 3D fabrication process for a wide range of materials,such as polymers, metals and ceramics [1].

Introduction1.1 Introduction to Additive Manufacturing143Additive manufacturing (AM) is a novel 3D fabrication process for a wide range of materials,such as polymers, metals and ceramics [1]. The fabrication by AM refers to processes in whichmaterial is added in layers guided by computer aided design (CAD) model to create a three-dimensional object [2]. There are distinct AM technologies including vat-photopolymerization,material extrusion, and binder jetting (BJ), just to mention few [3]. In the last decade, there hasbeen an increasing effort to characterize AM ceramic materials by using different techniques [3-5], namely with diverse spectroscopic techniques such as X-ray diffraction (XRD) spectroscopicand electron optical and surface chemical characterization techniques.

The recent advances insmall-scaled electronics, combined with computational equipment had allowed scanning electronmicroscopy (SEM) to see morphology and chemical analysis through energy dispersive X-rayspectroscopy (EDS) in specific areas of samples under investigation to study the surface offabricated components [2].1.2 Introduction to BJ processThe BJ process is a suitable technique for fabrication of 3D structures from precursor ceramicpowder materials. The BJ process uses two main materials; a powder material and a binder. Thebinder acts as an adhesive between the powder layers and it is dispensed selectively into thepowder bed through an inkjet print head that is able to displace along the "X" and "Y" axes of themachine. After printing, the binder can be cured, and a de-binding step carried out using theappropriate thermal post-processing [2].

After printing, the binder can be cured, and a de-binding step carried out using theappropriate thermal post-processing [2]. Several steps are incurred in BJ processing of ceramiccomponents: first, the ceramic powder material is spread over the build platform using a roller;second, the inkjet print head deposits the binder as adhesive on top of the powder where requiredby the design of the component; and third the build platform is lowered by a prescribed layerthickness, and another layer of ceramic powder is spread over the previous layer. Steps 1 through3 are repeated until the entire component has been printed. A schematic of the BJ process isshown in Figure 1.144Figure 1.

Steps 1 through3 are repeated until the entire component has been printed. A schematic of the BJ process isshown in Figure 1.144Figure 1. Schematic representation of BJ process1.3 AlN ceramic materialAlN has been used in the electronic industry since the past decade due to its combination ofremarkable properties, such as wide band gap, high thermal conductivity, low thermal expansion,high strength, optical transparency, and piezoelectric properties [4]. AlN is a covalently bondedceramic, it is synthesized from aluminum (13Al) and nitrogen (7N) in the stoichiometric aspresented in Table 1. It does not occur naturally, and once synthesized it is stable in inertatmospheres at temperatures over 2100 oC. After fabrication of ceramic materials, componentsare sintered and also subjected to hot isostatic pressing (HIPing) method to improvedensification.

After fabrication of ceramic materials, componentsare sintered and also subjected to hot isostatic pressing (HIPing) method to improvedensification. The HIPing method requires high pressures and temperatures intervals at differentcycles under inert atmosphere [5]. The ability to use BJ for the fabrication of AlN exhibitinghigh thermal conductivity and strong dielectric properties, highlights the potential for thefabrication of components or devices for future electronic packaging applications [3].Table 1. AlN stoichiometric chemical compositionIn this work, we report on the characterization of the thermal and other material properties forAlN produced using BJ, and highly densified through the use of post-fabrication heat treatments,for the first time. The components were fabricated using an ExOne M-lab equipment. Afterfabrication, components were sintered and also subjected to HIPing method to improvedensification.

The components were fabricated using an ExOne M-lab equipment. Afterfabrication, components were sintered and also subjected to HIPing method to improvedensification. After sintering and HIPing, various properties of the material, including thermal145conductivity, crystal structure and chemical composition, were analyzed using SEM, EDS, andXRD.2. Experimental Details2.1 Fabrication of components by BJIn this work, the AlN precursor powder was sourced from American Element, with a measuredpurity of 99.9%. The powder had an average particle size of ~45 μm. Fabrication of componentswas performed in an ExOne M-Lab (Irwin, PA serial number: 0600H2). Other fabricationparameters used included a binder saturation level of 200%, powder packing ratio of 57.62%,and a feed ratio to layer of 1.75. Several geometries were designed and fabricated to test theresolution capabilities of the BJ to process AlN.

Several geometries were designed and fabricated to test theresolution capabilities of the BJ to process AlN. The geometries included a set of tubes having adiameter of 1”, and varying thicknesses of 1/8”, 1/16”, 1/32”, and 1/64”; at two different heightsof 1” in and ½”, respectively. Other geometries built included pellets with a diameter of 1/2” andthickness of 0.11”. The flow diagram for BJ of AlN components fabrication process is shown inFigure 2. The flow diagram shows the followed BJ fabrication process which main stages gounder fabrication, densification and characterization of material properties.Figure 2. Fabrication flow diagram process for BJ of AlN components2.1.1 Curing, sintering and HIPing parameters146After printing, the AlN green bodies were subjected to a binder curing thermal process at 200 oCfor 2 h. After curing, a sintering cycle was performed in a Sentro Tech furnace 1600.

After curing, a sintering cycle was performed in a Sentro Tech furnace 1600. A firstsintering cycle was performed, at 1500 oC for 1h with and without vacuum conditions andwithout use of a carbon crucible in order to understand the densification and chemicalcomposition of fabricated components inside the furnace elements. . A second sintering cyclewas performed with vacuum and in an inert atmosphere and using a carbon crucible. Then, thefurnace was vacuum purged three times at -27 inHg, and a positive flow of argon (Ar) was usedto create an inert atmosphere. Then, the sintering cycle was set with a ramp rate of 5 oC/min,followed by three dwell ramps at 150 oC for 1 h, 600 oC for 20 min and at 1500 oC for 5 h. Thethree dwell ramps used achieved humidity removal, binder evaporation, and increased thestrength of AlN components, respectively. Finally, using an AIP-30H furnace model, a HIPingcycle was tested to increase the density of the fabricated components.

Finally, using an AIP-30H furnace model, a HIPingcycle was tested to increase the density of the fabricated components. The HIPing cycleconsisted of a ramp rate of 5°C/min at 1900 °C in a nitrogen atmosphere at 30,000 psi with an 8h hold. The sintering and HIPing profile used are presented in figure 3.Figure 3. Sintering and HIPing parameters used to densify AlN components.2.1.2 Shrinkage and density measurementsThe dimensional shrinkage analysis (DSA) was done using a Metrology Multisensor OGP with azoom up for accuracy of 10x magnification. The DSA of fabricated components with sinteringand HIPing required the measurements of components at exterior diameter (Dext), interiordiameter (Dint), and wall thickness (width), respectively.

The DSA of fabricated components with sinteringand HIPing required the measurements of components at exterior diameter (Dext), interiordiameter (Dint), and wall thickness (width), respectively. Density measurements of manufactured147components were calculated using the standard density formula where the mass is divided by itsvolume.2.2 SEM and EDSMorphology aspects and chemical analysis were performed by using a SEM from HITACHISU3500 system with an integrated EDS spectroscopy from OXFORD Instrument X-MaxN brand,respectively.2.3 XRDXRD was done using a PANalytical Empyrean system with a CuKα radiation source at 40 keVand 30 mA in the 2θ range between 20o and 80o, at 0.02o steps every 4 seconds.

The AlNcrystallographic indexation was done using #ICDD: 00-025-1133 for hexagonal (C6v4-P63mc).2.4 Thermal conductivity measurementsThe thermal conductivity (TC) measurements were performed using a Netzsch LFA 457MicroFlash table-top instrument that uses the laser flash method (LFM) for thermal diffusivitymeasurements. The LFM is capable of determining the specific heat profile by using graphite orany standard that is pre-loaded into the software. Because the software knows what to expectfrom the graphite standard, it has a reference point for comparison to determine the specific heatof the unknown sample. Two pellet components of AlN with a diameter of 0.5” and thickness of0.11” were employed for this experimental at sintering and HIPing parameters, respectively.3. Results and Discussion3.1 Fabrication of components by BJThe ExOne M-Lab was used to successfully produce various components out of AlN precursorpowders.

Results and Discussion3.1 Fabrication of components by BJThe ExOne M-Lab was used to successfully produce various components out of AlN precursorpowders. Tubes with the dimensions previously discussed were all successfully fabricated. No148particular issues were found for any of the produced components, even those that were tall andhad the smallest thickness. Figure 4 shows the BJ produced green body AlN tube with a 1”diameter, thickness of 1/64”, and height of 1”, after curing at a temperature of 200 oC for 2 h.Figure 4. Tube different views; a) and b) CAD design of tube with a 1” diameter, thickness of 1/64”, and height of 1”; c) Printed green body AlN tube after curing at 200 oC for 2 h.The fabrication of this component required 18h and 35 min, consisting of 260 deposited layers ofAlN powder (at a layer thickness of 45μm), a spread powder layer time of 20 mm/sec and using22 passes of binder in between layers.

It is worth mentioning that the AlN phase is easy to loseduring fabrication and sintering of components due to the covalent bond between Al and N easyto contaminate at temperatures between 1500 oC to 1850 oC [6, 7]. Figure 5a) shows the tubeafter the sintering thermal process under vacuum conditions. As it can be appreciated, the tubemaintains its grey color, which indicates that is related to the AlN phase. Figure 5b) shows thetube without vacuum conditions (open atmosphere). During sintering in an open atmosphere, acolor change can be observed, where the material turned from grey to white which is related to apotential nitradation (oxidation) that occurred during the sintering process without vacuumthermal conditions.149Figure 5. a) AlN tube component sintered under vacuum conditions; b) AlN tube sintered in airconditions turned into aluminum oxide (Al2O3).Furthermore, Figure 6a) shows the tube after the sintering thermal process for de-binding (brownbody).

As it can be appreciated, the tube turned to a black color, which was attributed to the useof a glassy carbon crucible to protect the AlN tube from potential oxidation or contaminationfrom the furnace heating elements during the sintering process.Figure 6. a) Tube after the sintering thermal process using a carbon crucible to avoid oxidationand b) tube after HIPing process showing the white-grayish related to AlN composition.Finally, the finished AlN tube was subjected to the HIPing cycle to increase its densification.Figure 6b) shows the tube after the HIPing cycle. The tube changed coloration from black(sintered tube sample) to white-grayish which is related to AlN ceramic phase [7].

The tube changed coloration from black(sintered tube sample) to white-grayish which is related to AlN ceramic phase [7]. It is worthmentioning that components were produced using BJ without sintering aids.3.1.1 Shrinkage and density measurements150The 1” diameter, 1/8” thickness tube component was measured using a SmartScope OGP tocalculate the shrinkage it experienced after the post-processing steps of sintering and HIPing.Figure 7 shows the image as seen from the SmartScopeOGP top view of a segment of HIPedtube. It shows good packaging of the layers, a good texture and well defined and uniform edgesaround the surface of the component.`Figure 7. a) Picture of a segment at top view of AlN tube after HIPing process.Table 2 shows the calculated shrinkage experienced by the AlN tube with nominal externaldiameter and wall thickness of 1”and 1/8”, respectively.

Compared with the nominal values inthe as fabricated condition, specimens that were sintered and subjected to the HIPing cycle hadoverall shrinkage values of: -0.025”, -0.001” and +0.111” for the external, internal diameters andwall thickness dimensions (width).Table 2. Tube with sintering and HIPing for dimensions analysis.The measurements taken with the SmartScopeOGP showed that there was shrinkage ofdimensions of the AlN 3D printed components which are in a complete agreement with HIPingprocess for sintering ceramics materials [6-8]. The reported density measured for this AlN tubeat various stages during the fabrication ranged from ~ 1.07 g/cm3 after curing at 200 oC, ~ 1.13g/cm3 after sintering at 1500 oC, and finally to ~ 1.96 g/cm3 after the HIPing cycle. The density151values mentioned correspond to ~ 32 %, ~ 35 %, and ~ 60 % of the theoretical density for AlN[7].3.2 SEM-EDSMicroscopy through SEM was employed to characterize the density of components.

One of thereasons densification is difficult, is because of trapped gases evolved from the powder surface orthe decomposition of binder trapped in closed pores [8]. It is well known that the post thermalprocessing for AlN is difficult to achieve due to its covalent nature [9].3.2.1 Sintering with and without vacuum conditionsThe fabricated components were analyzed by SEM and EDS, the first analysis was performedwith samples that were subjected to sintering process conditions with and without vacuum at1500 oC and no without use of a carbon crucible, were studied to understand the thermalbehavior of the material in different thermal conditions at the same temperature.Figure 8a) shows a representative SEM image at 2000X magnification from the AlN componentsintered at 1500 oC with vacuum and without crucible protection. Figure 8b) shows the EDScorresponding to the chemical analysis found in sintered component.

Figure 8b) shows the EDScorresponding to the chemical analysis found in sintered component. The percent content ofchemical elements found were 38%, 60% and 12%, for aluminum, oxygen and molybdenumrespectively. The traces of molybdenum (Mo) were attributed to contamination coming from thefurnace heating elements made having a chemical composition containing this element. It isworth mentioning that besides oxygen, AlN has a high affinity for Mo which inducescontamination [7].Figure 8. AlN sintering at 1500 oC at vacuum without crucible; a) SEM and b) EDS152Figure 9a) shows the SEM image of the component sintered at 1500 oC without vacuum and nocrucible covering. The image shows random sized AlN polycristals joined together withoutsintered particles properties. Figure 9b) shows the corresponding EDS (from the area indicated inFigure 8a). In the image, it can be seen that chemical elements Al, O and Fe were found at 52%,47% and 3% content respectively.

Figure 9b) shows the corresponding EDS (from the area indicated inFigure 8a). In the image, it can be seen that chemical elements Al, O and Fe were found at 52%,47% and 3% content respectively. This confirmed that the components would have undergoneoxidation. Figure 9c) and 9d) show EDS maps showing profiles in green aluminum and red foroxygen.Figure 9. AlN sintering at 1500 oC in air atmosphere.EDS results for both sintered samples with and without the use of vacuum and no cruciblecovering, showed that contamination was present in components as exhibited by the presence ofMo and Fe, which were related to the sintering furnace heating elements.

As mentioned beforeoxygen is well known combining to AlN, it was observed a potential presence of oxygen withinthe components with and without vacuum of 60% and 47%, respectively.3.2.2 Sintering under Ar atmosphere and HIPingFurthermore, sintering conditions were explored using the vacuum and argon (Ar) inertatmosphere and with a carbon crucible to protect components from furnace chamber153contamination. Figure10 from a) to d) shows a set of SEM images at different magnifications ofthe morphology structure profile of the component subjected to a sintering thermal treatment at1500oC in Ar atmosphere and vacuum conditions. This image indicates agglomerations ofparticles and cluster of smaller microparticles bound together due to the binder, the compactioncreated by the roller during printing, and the sintering conditions.Figure 10.

From a) to d) SEM images of AlN at different magnifications with sintering at 1500 oCin Ar atmosphere/vacuum and covering with a carbon crucible.The regions indicated in blue squares in Figure 10b) correspond to the regions that were used forEDS chemical analysis composition. The EDS results were found to consist of 65% Al, 15% N,and 15% C, respectively. The occurrence of carbon was traced to the use of a glassy carboncrucible in order to protect the components being sintered from the heating elements. Thecrucible and inert gas atmosphere aided in preventing metallic contamination traces from theinterior of the furnace and avoid oxidation of material.Figure 11a) shows a representative SEM image at the cross section of the AlN tube that wassubjected to the HIPing process at 1900 oC. This view allows observation of homogeneoussintered microparticles exhibiting coalescence with well-formed boundaries.

This view allows observation of homogeneoussintered microparticles exhibiting coalescence with well-formed boundaries. The regionsindicated in white squares were used for EDS chemical analysis composition results.154Figure 11. a) SEM of AlN with HIPing cycle at 1900 oC and b) EDS results.Figure 11b) corresponds to the EDS results, showing that the components subjected to theHIPing cycle had a chemical trace corresponding to 65% Al, 32% N, and 3% O. This chemicalcomposition obtained belongs to original atomic/mass of Al and N in AlN stoichiometriccomposition. The oxygen trace can be the result of trapped porous structures due to thefabrication process.3.3 XRDThe formation of pure crystalline AlN phase by HIPing cycle was confirmed by XRD. Figure12shows the XRD results of AlN printed components after the HIPing process at 1900 oC.

Figure12shows the XRD results of AlN printed components after the HIPing process at 1900 oC. TheXRD pattern for BJ printed AlN component corresponds to wurtzite crystal structure showingthe principal plane directions at (001), (002) and (101) which is in a complete agreement withthose reported by Xiong et al., [9].155Figure 12. XRD of AlN components after HIPing process.3.4 TC measurementsHere we present the TC properties results measured for an AlN pellet component fabricated byBJ according to the previously discussed dimensions and sintering thermal conditions.3.4.1 TC at sintering conditionsFigure 13 from a) to c) shows the TC measurements of AlN 3D printed pellet with sinteringprocess from 25 oC to 500 oC, results are presented such as a) specific heat, b) thermaldiffusivity, and c) TC, respectively. TC results were found at 2.8 W/m*K and 1.4 W/m*K atroom temperature and 500 oC, respectively.156Figure 13.

TC results were found at 2.8 W/m*K and 1.4 W/m*K atroom temperature and 500 oC, respectively.156Figure 13. TC measurements of AlN 3D fabricated pellet with sintering process from 25 oC to 500 oC: a) Specific heat (J/gK), b) Thermal diffusivity (mm2/s), and c) TC (W/mK).3.4.2 TC at HIPing conditionsThe TC properties results of AlN pellet component with HIPing conditions are presented infigure 14 from a) to b). Fig. 14a) shows a plot showing the specific heat measurements obtainedin the temperature range from 25 oC to 500 oC. Figure 14b) shows a plot with the thermaldiffusivity values in the same range of temperatures. Figure 14c) shows a plot of TC as afunction of the temperature in the range from 25 oC to 500 oC, in the range from 4.82 W/m*K to3.17 W/m*K.Figure 14.

Figure 14c) shows a plot of TC as afunction of the temperature in the range from 25 oC to 500 oC, in the range from 4.82 W/m*K to3.17 W/m*K.Figure 14. TC measurements of AlN 3D printed pellet with HIPing from 25 oC to 500 oC: a) Specific heat (J/gK), b) Thermal diffusivity (mm2/s), and c) TC (W/mK).The TC results obtained are mainly affected by the average size of microcrystals, density, andsynthesis method and fabrication process of components [9]. The fabrication method employedhere of BJ with HIPing obtained good printed components. Nevertheless, AlN 3D printedcomponents densification through sintering and HIPing may induce some contamination andoxidation. It is well known that oxygen is a typical contamination of AlN, located at the surfaceof AlN grains as aluminum oxide or oxynitrides [10].

It is well known that oxygen is a typical contamination of AlN, located at the surfaceof AlN grains as aluminum oxide or oxynitrides [10]. In single crystals of high purity this valuecan reach ~ 320 W/(m*K) at room temperature, but for sintered polycristals values in the rangefrom 17-285 W/(m*K) are more typical. In general, AlN 3D printing of polycrystalline ceramicsexhibit lower thermal conductivities compared with their associated AlN single crystals [11]. Asa result of impurities or lattice defects in AlN 3D fabricated components, phonon scattering mayoccur which leads to a decrease in thermal conductivity. One of the differences between AlN157single crystals, polycrystalline ceramic and AlN 3D fabricated components is that, forpolycrystalline materials, the heat carrying process is by mean of phonons which propagatedifferently in each system network [12].

In single crystals the heat carrying is influenced by theirnucleation, stoichiometric chemical homogeneity, mechanical stresses and phase transition in TCceramics materials.4. ConclusionsIn this work we have presented results of BJ production of AlN components having high qualitythermal conductivity properties for the first time. The structural aspects and thermal conductivityproperties were analyzed by using LFM, SEM, EDS and XRD, respectively. The XRD patternresult showed the characteristic crystallographic planes of AlN 3D printed components at (001),(002), and (101) related to wurtzite crystal structure. The SEM results showed that the BJproduced AlN components subjected to HIPing exhibited a crystalline structure due to theformation of well-defined grain boundaries. With LFM, it was found that thermal conductivityproperties of HIPed BJ produced AlN in the interval from 25 oC to 500 oC ranged from 4.82W/m*K to 3.17 W/m*K, respectively.

With LFM, it was found that thermal conductivityproperties of HIPed BJ produced AlN in the interval from 25 oC to 500 oC ranged from 4.82W/m*K to 3.17 W/m*K, respectively. In summary, the observed values of the printingparameters, binder density, powder packing rate, binder saturation level, and thermalconductivity properties found in BJ produced AlN components were shown be useful for thedevelopment of new devices for use in electronic devices including heat sinks. In futureexperiments, we propose sintering of BJ AlN at higher temperatures with the prospect ofincreasing densification. Future work also includes the application of BJ to other technical orengineering ceramics.AcknowledgmentsCarlos A. Diaz Moreno thanks Mexico’s Consejo Nacional de Ciencia y Tecnología(CONACYT) for support through Postdoctoral Abroad Program, Solicitation No. 250381 at TheUniversity of Texas at El Paso. Carlos Diaz special thanks to J.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceBINDER SATURATION, LAYER THICKNESS, DRYING TIME AND THEIR EFFECTS ON DIMENSIONAL TOLERANCE AND DENSITY OF COBALT CHROME - TRICALCIUM PHOSPHATE BIOCOMPOSITEJohn Ruprecht, Kuldeep Agarwal and Shaheen AhmedDepartment of Automotive and Manufacturing Engineering Technology, Minnesota State University, Mankato, 56001, USAbstractTraditional metals such as stainless steel, titanium and cobalt chrome are used in biomedical applications (implants, scaffolds etc.) but suffer from issues such as osseointegration and compatibility with existing bone. One way to improve traditional biomaterials is to incorporate ceramics with these metals so that their mechanical properties can be similar to cortical bones. Tricalcium phosphate is such a ceramic with properties so that it can be used in human body.

Tricalcium phosphate is such a ceramic with properties so that it can be used in human body. This research explores the use of binder jetting based additive manufacturing process to create a novel biocomposite made of cobalt chrome and tricalcium phosphate. Experiments were conducted and processing parameters were varied to study their effect on the printing of this biocomposite. Layer thickness, binder saturation and drying time affected the dimensional tolerance and the density of the green samples. This effect is important to understand so that the material can be optimized for use in specific applications.IntroductionBiomaterial is any material, natural or man-made, that can perform a body function or can replace a body part or tissue.

Depending on their application, biomaterials can be made from polymers, metals, ceramics or composites [1-3].Biomedical implants can be classified into 3 major categories: External to the body (non- clinical, which includes surgical instruments, prosthetics etc.), Internal to the body & permanent (includes hip implants, knee implants, stents etc.) and Internal to the body & temporary (includes scaffolds, degradable screws and drug delivery systems) [1-3, 14].According to FDA, a ‘‘permanently implantable device is a device that is intended to be placed into a surgically or naturally formed cavity of the human body for more than one year to continuously assist, restore, or replace the function of an organ system or structure of the human body throughout the useful life of the device.’’ Some examples include knee and hip implants.

Temporary implants are commonly used in sports surgeries, such as in shoulder and knee ligamentous reconstruction and spinal reconstructive surgery [5, 6, 14].The most common biomaterials for implants are metals and alloys, ceramics and polymers. In the past few years, Cobalt alloys have gained popularity for being used as biomaterials for implants. Bones have elastic moduli 7 -30 GPa, yield stress 30-70 MPa, compressive strength 100-148230 MPa, and tensile strength 70 – 150 MPa. First generation of implants focused on replacement of the bone with a metal implant. However, mechanical properties of metals differ considerably from natural bone: Cobalt chrome alloys elastic modulus = 210 GPa, yield strength = 120 - 600 MPa, tensile strength 190 - 800 MPa [3, 4, 7].These differences lead to stress shielding resulting in loosening of the implant due to degradation of human tissues around implants and, consequently, further surgeries to replace the implants.

Ceramics are inorganic materials with high compressive strength and biological inertness that make them suitable for scaffolds used in strengthening or replacing damaged bones and tissues. The most commonly used bioceramics are metallic oxides (e.g., Al2O3, MgO), calcium phosphate (e.g., hydroxyapatite (HA), tricalcium phosphate (TCP), and octacalcium phosphate (OCP)), and glass ceramics (e.g. Bioglass, Ceravital).Calcium phosphates have the best biocompatibility and properties closest to natural bones: elastic modulus 7-13 GPa, compressive strength 350-450 MPa, tensile strength 38-48 MPa and flexural strength 100 -120 MPa. However, they have poor fracture toughness and tensile strength that limits their application to bioimplants. Several in vitro and vivo works have shown that calcium phosphates support the adhesion, differentiation and proliferation of osseogenesis-related cells (e.g., osteoblasts, mesenchymal stem cells), besides inducing gene expression in bone cells.

The most important calcium phosphate is hydroxyapatite (HA, Ca10(PO4)6(OH)2) with chemical characteristics similar to hard tissues such as bone and teeth, that promotes hard tissue ingrowth and osseointegration when implanted into the human body. The porous structure of this material can be tailored to suit the interfacial surfaces of the implant. As a bulk material, HA lacks sufficient tensile strength and is too brittle to be used in most load bearing applications. In such cases, HA is coated onto a metal core or incorporated into polymers as composites. The ceramic coating on the titanium implants improves the surface bioactivity but often fails as a result of poor ceramic/metal interface bonding [8-11].(cid:2745)-TCP and (cid:533)-TCP are the two crystalline varieties of HA of interest in biological applications. (cid:533)-TCP is the thermodynamically stable form at low temperature. It transforms into (cid:2745)-TCP in the temperature range 1120–1170 oC.

(cid:533)-TCP is the thermodynamically stable form at low temperature. It transforms into (cid:2745)-TCP in the temperature range 1120–1170 oC. (cid:533)-TCP is generally preferred in sintered ceramic implants, while (cid:2745)-TCP is more commonly used in bone graft cements because of its hydrolysis properties. The requirements that allow bone ingrowth are a porosity of 30–70 vol % and a pore diameter between 300 and 800 μm, mechanical properties of 0.5–15 MPa similar to cancellous bone.Thus there is a drive in the biomedical industry to create novel materials that behave very similar to bone and can be used for multiple applications: from permanent to temporary implants. Furthermore, these materials need to be manufactured in a manner that would create porosity in situ for biological applications.Due to the versatility of additive manufacturing, it is gaining a lot of popularity in the field of bone implants.

Selective Laser Sintering, Selective Laser Melting, Electron Beam Melting and Binder jet manufacturing have all been used to create various porous structures for biomedical implants [7, 9]. To accomplish the various different requirements of the implants it is necessary to create biocomposites that can have the strength properties of metals as well as the biological properties of bioceramics.149This research explores the use of binder jetting based additive manufacturing process to create a novel biocomposite made of cobalt chrome (CC) and tricalcium phosphate (TCP). Experiments were conducted and processing parameters were varied to study their effect on the printing of this biocomposite. Layer thickness, binder saturation and drying time affected the dimensional tolerance and the density of the green samples.

Layer thickness, binder saturation and drying time affected the dimensional tolerance and the density of the green samples. This effect is important to understand so that the material can be optimized for use in specific applications.Binder Jet-based Additive ManufacturingBinder Jetting AM is a Drop on Demand (DoD) inkjet printing process in which binder is emitted through a nozzle to form a short jet. This jet condenses into a drop and the position at which each drop lands on substrate is controlled by relative motion between drop and the substrate. The nozzle head is piezoelectric and uses the deformation of ceramic element to generate a pressure pulse needed to eject the binder.

The nozzle head is piezoelectric and uses the deformation of ceramic element to generate a pressure pulse needed to eject the binder. Typical drop diameters vary from 10 μμm – 100 μm, drop volumes vary from 0.5 – 500 pl and the drop speeds are 5-8 m/s [15].The main technique of manufacture using the Binder Jet process is as follows: (a) The CAD file is sliced into layers and a STL file is generated, (b) Each layer begins with a thin distribution of powder spread over the surface of a powder bed, (c) Using a technology similar to ink-jet printing, a binder material selectively joins particles where the object is to be formed, (d) A piston that supports the powder bed and the part-in-progress lowers so that the next powder layer can be spread and selectively joined, (e) This layer-by-layer process repeats until the part is completed, (f) Following a heat treatment, unbound powder is removed and the metal powder is sintered together. Fig.

Fig. 1 shows the details of the whole process.dFig. 1: Schematic of the Binder Jet Process (Courtesy The ExOne Company)Process parametersThe Binder Jet process described above can be divided into 3 basic steps: 1) Binding, 2) Curing and 3) Sintering. There are various process parameters that can be changed to obtain a customized part in each of these steps. These include powder size, layer thickness during binding, part150orientation in bed, heater power, roller speed, curing temperature, curing time, sintering time, sintering temperature, and sintering atmosphere.

The images are shown in Fig. 2.Co 212-H was used as a benchmark and is called Sample Set 1. The two powders weremixed in with 80% Co 212-H -20% TCP volume fraction and this is called sample set 2.SEI MSU10kVWD14mmSS4Fig. 2: SEM Micrographs for Co 212-H and TCPThe 2 sample sets were then used as the input powders in the binder jet additive manufacturing. To understand the effect of printing parameters of the binder jet process on the print quality, multiple factors and levels were chosen based on prior experience (Table 2). After the printing, all the samples were cured at 175 oC for 3 hours.The print quality was measured by 2 means: dimensional tolerance compared to the CAD model and the density of the part after the printing. Two different size of parts were printed to study the effect of size on the dimensional tolerance. The CAD drawings of the parts with nominal dimensions is shown in Fig. 3.

Two different size of parts were printed to study the effect of size on the dimensional tolerance. The CAD drawings of the parts with nominal dimensions is shown in Fig. 3. Eight parts of 10mm and 3 parts of 30mm were printed for the corresponding experimental run.151S.No. 1 2 3 4Factor Layer Thickness Binder Saturation (as % of printhead max) Drying Time between layers Sample Size (Outside dimension) Table 2: Factors and levels for Design of experimentsLevels 60 μμm, 90 μμm 80, 100 15 sec, 20 sec 10 mm, 30 mmIf a full factorial experimental design would be done on these parameters, it would require a total of 16 experiments. Instead a half factorial experimental design was chosen and 8 experiments were conducted. The experimental design matrix is shown in Table 3.Fig. 3: CAD models with dimensions for 10 mm and 30 mm samplesExperiment No.

The experimental design matrix is shown in Table 3.Fig. 3: CAD models with dimensions for 10 mm and 30 mm samplesExperiment No. 1 2 3 4 5 6 7 8Layer Thickness Binder SaturationDrying Time 15 sec 20 sec 20 sec 20 sec 15 sec 15 sec 15 sec 20 sec90 μμm 60 μμm 90 μμm 90 μμm 60 μμm 90 μμm 60 μμm 60 μμm100 % 100 % 100 % 80 % 80 % 80 % 100 % 80 % Table 3: Experimental design matrixSample Size 10 mm 10 mm 30 mm 10 mm 10 mm 30 mm 30 mm 30 mmResultsAfter the parts were printed, they were weighed and measured for the outside X, outside Y, inside X and inside Y dimensions.

6: Dimensional variation along 2 directions for sample set 2 (80% Co 212-H, 20% TCP)The samples from various experiments were also photographed to look at the printing errors such as bleed out, distorted walls, rough surface, excess material etc. These are shown in Fig. 7-8.(b) Fig. 7: (a) 10mm samples with printing bleed out, (b) 10mm sample with printing error (c) Good 10mm samples(a)(c)(a) (b) Fig. 8: (a) 30mm sample depicting printing error (b) Good 30mm sampleThe response surface analysis of all the results was done using MINITAB. Only the main factors were considered in the analysis and the interaction between the factors was ignored for this study since the aim is to find the process parameters imports for dimensional variations. A typical fitted response surface is shown in Fig.

A typical fitted response surface is shown in Fig. 9.154■ Expt 1■ Expt2■ Expt3■ Expt4■ ExptS■ Expt6■ Expt7■ Expt80%HAResponse Surface Regression: Outside X versus LT, BS, DT, SSlysis of Variana?Linear LT BSDT ssDF Mj SS Adj MS 502. 72 70719.03 7 3519.03 879.74 123756.20 4 3518.97 4.39 0.03 1 14.67 0.10 1 0.34 0.00 1 1 3518.83 3518 .83 495005.41F - Value P- Value 0. 000 0. 000 0. 043 0. 000 0.566 0. 0000.03 0.10 0.0036 0.26 43 3519.280.01S 0.0 843129 99. 99%-R- sq R- sq (adj) R - sq(pred) 99. 99%-99. 99%-Fig. 9: Response surface result from MINITAB. LT: Layer Thickness, BS: Binder Saturation, DT: Drying time and SS: Sample SizeAll the response surfaces were analyzed and the results of the P-values were summarized to identify the significant factors in each case.

Results from the various tests, the box plots and the response surface analysis show the following:155a) There is a difference in the dimensional variations and density for both the 0% and 20%TCP samples.b) Layer thickness and sample sizes are significant factors in both the sample sets. Thus, the dimension of the part, its surface and density are significantly affected by the layer thickness of the binder jet process. This is expected because the layer thickness affects the compaction of each layer and hence determines how much binder can saturate each powder layer.c) Drying time is not a significant factor in 0% TCP samples but is important for the 20% TCP samples. The TCP is a much softer material than Co 212-H and hence the binder saturates the powder a lot more in this case.

The TCP is a much softer material than Co 212-H and hence the binder saturates the powder a lot more in this case. This can result in material bleed outs or layers being dragged on another layer if it has not dried enough.d) Similarly, binder saturation affects the dimensional accuracy more in case of 20% TCP samples than in the case of 0% TCP samples. In fact, in 0% TCP samples, binder saturation does not affect the inside dimensions. This is because once a layer is formed, the outside surface is more prone to the distortion or dislocation than the inside surface.e) The 30 mm samples have much smaller variation in dimensions than the 10 mm samples. This might be because the larger surface area is more uniformly spread and dried than the smaller surface area of 10 mm samples.f) By looking at the samples visually and their dimensions, Experiment 5 (60 μμm Layer thickness, 80% binder saturation and 15 sec drying time) performed the best for 0% TCP Samples, 10mm size.

In case of 20% TCP samples, experiment 4 (80 μm Layer thickness, 80% binder saturation and 20 sec drying time) performed the best.g) There is a wide variation in the density of green parts depending upon the process parameters. This property can be taken advantage of to control the porosity of the parts made by various materials.ConclusionsTraditional metals such as stainless steel, titanium and cobalt chrome are used in biomedical applications (implants, scaffolds etc.) but suffer from issues such as osseointegration and compatibility with existing bone. One way to improve traditional biomaterials is to incorporate ceramics with these metals so that their mechanical properties can be similar to cortical bones. TCP is such a ceramic with properties so that it can be used in human body. This research explores the use of binder jetting based additive manufacturing process to create a novel biocomposite made of cobalt chrome and tricalcium phosphate.

This research explores the use of binder jetting based additive manufacturing process to create a novel biocomposite made of cobalt chrome and tricalcium phosphate. Experiments were conducted and processing parameters were varied to study their effect on the printing of this biocomposite.It is found that layer thickness is significant factor for printing with 0% TCP and 20% TCP samples. However, drying time is significant for 20% TCP samples because of the soft nature of TCP powder. Similar effect is found with binder saturation. Thus, printing parameters are crucial in manufacturing parts from binder jet additive manufacturing.156Further research needs to be conducted to take the best printing parameters and use the samples from this configuration and study the effect of sintering parameters on the strength and biocompatibility of these samples.References1. Tarafder, S., 2013.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperBinderless Jetting: Additive Manufacturing of metal parts via jetting nanoparticlesYun Bai and Christopher B. WilliamsDesign, Research, and Education for Additive Manufacturing Systems Laboratory Department of Mechanical Engineering Virginia TechAbstractBinder Jetting AM has been used to fabricate metal parts by first jetting a binder into powder bed; the resulting green part is then thermally post-processed wherein the binder is removed and the metal particles are sintered. In this work, the authors replace conventionally- used polymeric binders with nanoparticle suspensions as a means for binding metal powder bed particles together.

In this work, the authors replace conventionally- used polymeric binders with nanoparticle suspensions as a means for binding metal powder bed particles together. After being deposited into the powder particles’ interstices, the jetted nanoparticles are sintered at a low temperature via a heated powder bed to provide strength to the printed green part. Regions of the powder bed that do not receive the jetted nanoparticle suspension remain as loose powder as the sintering temperature of the nanoparticles is significantly lower than the larger powder bed particles. The concept of printing metal by jetting a nanoparticle binder made of the same material is demonstrated in the context of copper through printing copper parts with satisfactory green strength.Key words: Binder Jetting, Nanoparticles, Inkjet Printing, Sintering, Copper1.

Introduction1.1 Binder Jetting overviewThe Binder Jetting Additive Manufacturing (AM) process can be used to fabricate metalparts by selectively inkjet printing a liquid binding agent into a powder bed, followed by post- process sintering of the printed green part. As shown in Figure 1, the jetted binder droplets interact with the powder particles to form primitives that stitch together to form a cross-sectional layer. Once a layer is printed and thermally dried/cured, a new layer of powder is recoated on top of the previous layer which is then printed and stitched to the previous layer by the jetted binder. The layer-by-layer process is repeated to create the complete green part. The unbound loose powder in the bed that surrounds the part supports overhanging structures during the build, and can be removed after printing via compressed air.

The unbound loose powder in the bed that surrounds the part supports overhanging structures during the build, and can be removed after printing via compressed air. Once depowdered, the green part is placed in a furnace to burn off the binder and to sinter the powder particles together to obtain final density and strength.1249Figure 1. Green part printing process in Binder JettingBinder Jetting of metals differs from direct-metal AM technologies in that it separates thegeometry formation from the thermal processing. As Binder Jetting of metals functionally separates part creation from powder sintering, the requirement of anchors and/or heat sinks due to the residual stresses imposed by large thermal gradient in direct-metal AM processes is avoided.

The ability to fabricate a part in a powder bed without the need for built anchors enables Binder Jetting to create large and geometrically complex parts without the need for the difficult and laborious support/anchor removal in post-processing printed metal parts. In addition, Binder Jetting is an inherently scalable and cost-effective technology, which has been demonstrated in many commercial systems (e.g., Desktop Metal’s Single Pass Jetting technology is 100x faster and 20x lower cost than laser-based systems1).1.2 Binder selectionThe selection of binder material for Binder Jetting is critical as it determines the success of creating satisfactory green parts and affects the final properties of the sintered parts. Different types of binders have been used in Binder Jetting, including various solvents, colloids, and polymers (Utela et al., 2010).

Different types of binders have been used in Binder Jetting, including various solvents, colloids, and polymers (Utela et al., 2010). In modern commercial metal Binder Jetting systems, solvent-based binders that contain thermosetting polymers are commonly used to print green parts; thermal curing via an overhead (or integrated) heater removes the solvent and crosslinks the jetted polymer.

While polymer binders have demonstrated success with many metal systems, its use in Binder Jetting could add complexity to the manufacturing process and adversely affect the final product performance metrics:(i)(ii)(iii)the debinding of polymer binders typically require refined sintering profile to facilitate polymer pyrolysis and degassing, the green part could lose structural integrity (i.e., warping, creep) during sintering if the onset sintering temperature of powders is higher than the binder burn-out temperature, and the pyrolysis of polymer binder could leave residual ash, which could affect the purity (and thus mechanical, optical, electrical/thermal properties) of the final part.1.3 Nanoparticle binderIn this work, the authors experimentally investigate nanoparticle suspensions as analternative to replace polymer binders.

Theories and microscopic observations have shown that many metal nanoparticles manifest size-dependent meting behaviors and lower melting/sintering temperatures compared with the bulk materials (Asoro et al., 2009). For example, copper nanoparticles can sinter and neck within temperature ranges (150-300 °C) that are significantly1 https://www.desktopmetal.com/products/production/2250reduced from sintering temperature for micron-sized powders (Yu et al., 2011).

This unique feature of nanoparticles (sintering temperature reduction), when nanoparticles are jetted as a binder in Binder Jetting, can enable the selective cohesion of the powder bed when heat is supplied at each layer to sinter jetted nanoparticles while not affecting the powder in unprinted regions.Experiments are conducted in the context of copper, wherein copper nanoparticles are jetted into the powder bed, and sintered in the powder bed via an overhead heater, providing strength to the green part and acting as a permanent and structural binder in the final printed copper product. The low sintering temperature of copper nanoparticles ensures that only moderate heat is supplied to the powder bed, which ensures that the printed green part can be retrieved from the surrounding loose powder.Section 2 provides a critical analysis of the prior art, and establishes the rationale for thisapproach.

In general, these approaches were shown to improve material properties, but were still using an existing polymeric binder and relying on the curing of the polymer to provide green strength.The concept of using metal and ceramic particle slurries or colloids as a binder in Binder Jetting has been explored in prior work. Yoo et al. used jetted particulate suspensions as binders in steel powders in order to reduce the shrinkage by eliminating the need to sinter steel powders (Yoo, 1995). Carbonyl iron (1.7 µm) was suspended into water and inkjet printed to stainless steel powders (60 µm) followed by firing the entire powder bed in forming gas at 600-700 °C, under which temperature the interstitial carbonyl iron particles sintered without sintering the base powder. The printed part was “fairly weak” according to the authors, but was shown to be successfully retrieved from the powder bed and depowered. Hadjiloucas et al.

The printed part was “fairly weak” according to the authors, but was shown to be successfully retrieved from the powder bed and depowered. Hadjiloucas et al. used silver suspensions as the binder for printing molybdenum powders (Hadjiloucas, 1997). The printed interstitial silver particles melted and bonded the base powder particles after heating at 1000 °C, at which the base powder remained loose and ensured part retrieval from the powder bed. Techapiesancharoenkij et al. printed iron and nickel oxide slurries into Fe-30Ni base powders for printing bimetallic bars (Techapiesancharoenkij et al., 2004).

Techapiesancharoenkij et al. printed iron and nickel oxide slurries into Fe-30Ni base powders for printing bimetallic bars (Techapiesancharoenkij et al., 2004). After heat treatment at 400 °C in a reducing atmosphere, the printed bars can be retrieved from unbound powders.3251While the feasibility of using metal or ceramic particle suspensions as binder has beendemonstrated in the literature, the success of printing and material system selection is contingent upon the large sintering temperature gap between the base powder material and the binder material. As a result, dissimilar material systems composed of refractory powders and low- melting-point-material binders have been used with the requirement of sintering the powder bed in a controlled atmosphere furnace after printing. The feasibility of using a metal binder of the same material as powder in a single material system has not been demonstrated.3.

The feasibility of using a metal binder of the same material as powder in a single material system has not been demonstrated.3. Experimental methodsThe experiments were conducted to (i) determine the process parameters and establish the manufacturing process for fabricating copper parts via jetting copper nanosuspension as a binder, and (ii) explore the nanoparticle binder’s effect on green strength and green/sintered densities. To achieve this goal, the drying and sintering condition for printing copper nanosuspensions was first determined; then, the nanosuspension was printed as a binder for copper powders with different binder saturation ratios. Green part strength/density and sintered density were measured for both nanoparticle and polymer binder printed parts with different binder saturation ratios.The overall copper part manufacturing process chain followed an established procedurein authors’ prior work (Bai et al., 2017; Bai and Williams, 2015).

The material system and manufacturing process used for evaluating copper nanoparticle as binders is discussed in Section 3.1-3.3, with the characterization methods for printed copper green and sintered parts provided in Section 3.4.3.1 Material systemA gas atomized copper powder with median particle size of 17 µm and packing density of52% was chosen as the base powder material. A commercially available copper nanoparticle suspension with a solid loading of 20-26 w.t.% and average particle size under 100 nm was used (Gwent Electronic Materials Ltd.).

A commercially available copper nanoparticle suspension with a solid loading of 20-26 w.t.% and average particle size under 100 nm was used (Gwent Electronic Materials Ltd.). This product is designed for the electronics printing industry and has a proven rheology (30mN/m surface tension and 13 cps viscosity) that is suitable for inkjetting by commercial piezoelectric printhead.ExOne’s solvent-based polymer binder was selected to print parts as control group for theexperiments.3.2 Nanoparticle binder drying conditionFinding an appropriate drying condition is critical to ensure the success of printing a new binder. The polymer binder printed parts were dried following the standard procedure (a ceramic overhead heater set to 200 °C scans the printed layer at a speed of 5mm/s following the jetting pass).

The polymer binder printed parts were dried following the standard procedure (a ceramic overhead heater set to 200 °C scans the printed layer at a speed of 5mm/s following the jetting pass). As a result of this drying pass, the temperature of the powder bed surface reached to 170 °C when the heater is directly placed above (measured by an infrared thermometer).4252The drying condition (powder bed temperature and drying time) for nanosuspension was determined through experimentation based on the thermogravimetric analysis (TGA) performed on a TA Instruments platform.3.3 Manufacturing processCopper green parts were printed with copper nanosuspension or polymer binder on anExOne R2 3D printer. Powder was spread with 5mm/s with a feed-to-build ratio of 1.2:1. A standard 100 µm layer thickness was applied to the printing of both binders.The drop volume for each ink was measured before and after printing to ensure accuratebinder saturation ratio.

A standard 100 µm layer thickness was applied to the printing of both binders.The drop volume for each ink was measured before and after printing to ensure accuratebinder saturation ratio. While the drop volume measured for the standard polymer binder at a standard jetting waveform is consitent (101 pL), the drop volume measured for copper nanosuspension was inconsistent and varied in a range (65-75 pL). The binder saturation ratio for the nanosuspension is therefore selected by assuming a drop volume of 70 pL. A summary of measured drop volume and adopted binder saturation ratio is provided in Table 1.Table 1. Measured drop volume and adopted binder saturation ratioPolymer binder Copper nanosuspensionMeasured drop volume 101 pL 70 pLBinder saturation ratio 100% 70% & 140%The green parts printed by polymer binder or nanosuspension were retrieved from thepowder bed immediately after the printing finished.

The polymer printed parts were subjected to an additional post-process curing process (200 °C for 1 hour) to fully set the polymer binder, as prescribed by the standard procedure. Post-process curing was skipped for nanosuspension printed parts. The depowdered green parts were then sintered in a hydrogen furnace with pure hydrogen atmosphere, following an established sintering profile which features 5 °C/min heating to a 2-hour isotherm at 1075 °C.3.4 Green and sintered part characterizationGreen part strength was evaluated by identifying the peak load at specimen rupture in a 3-point bending test (MTS platform;1kN load cell), using ASTM B312 standard as a reference. In order to reduce the material and time consumption in making the specimen, the green strength specimen geometry was scaled down to 60% of the original size specified in the standard.

In order to reduce the material and time consumption in making the specimen, the green strength specimen geometry was scaled down to 60% of the original size specified in the standard. In addition, a constant displacement rate (0.1 mm/s) was used instead of controlled loading rate due to the instrument used.A series of 18mm×6mm×3mm test coupons were printed for evaluating green andsintered densities. Green part density was calculated from the measured part mass and dimensions. The sintered density was measured by oil impregnation and immersion method using Archimedes principle following ASTM Standard 962. A LEO (Zeiss) 1550 field-emission SEM was used to evaluate printed green parts. Three green strength specimens and three density coupons were printed and tested for each binder and saturation ratio.52534.

Three green strength specimens and three density coupons were printed and tested for each binder and saturation ratio.52534. Results4.1 Nanoparticle binder drying conditionFigure 2 shows a failed printing with the nanosuspension ink when using the polymerbinder’s standard layer drying condition (Section 3.2). In this failure, the recoating of an insufficiently dried layer printed by the copper nanosuspension has created an uneven powder bed surface and damaged the previous layer. Therefore, a sufficient drying and setting of the jetted binder is required to enable the printed layer with enough integrity that can withstand the compressive force imposed by the powder-spreading roller.Figure 2.

Powder bed defects occurred on top of the printed layer after recoating, caused by insufficient binder dryingAs shown in Figure 3, the complete drying of the copper nanosuspension requires higher temperature (200 °C) and longer time (10 minutes) than that of the polymeric binder when a 20 °C/min heating ramp was used for TGA. Through a series of experimental printing trials, it was determined that directly placing the overhead heater above the powder bed for 2 minutes maintains the powder bed surface temperature at 260 °C, which is capable of sufficiently drying the printed layer without incurring recoating defects. This drying condition also ensures leaving the copper powder bed unsintered and free-flowing, which enables printed green part retrieval (Figure 4b).10080%t h g i e W6040Nanosuspension20Polymer binder00100200300Temperature (°C)Figure 3. Thermogravimetric analysis (TGA) of the copper nanosuspension and polymer binder6254Figure 4.

Thermogravimetric analysis (TGA) of the copper nanosuspension and polymer binder6254Figure 4. (a) printed layer by copper nanosuspension and (b) printed green parts surrounded by brushed-off loose powder prior to retrieving from the powder bed4.2 Nanoparticle bondingThe source of the interparticular bonding in printed green parts is revealed in Figure 5-7.In polymer binder printed parts (Figure 5), cured polymers existed between copper particles, whereas in copper nanosuspension printed parts the powder particles are connected by the same material-copper (Figure 6).Figure 5. The surface of the green parts printed by polymer binder (100% saturation), characterized by SEM at magnification of 2kx (left) and 5kx (right)7255Figure 6.

The surface of the green parts printed by polymer binder (100% saturation), characterized by SEM at magnification of 2kx (left) and 5kx (right)7255Figure 6. The surface of the green parts printed by nanosuspension (70% saturation), characterized by SEM at magnification of 2kx (left) and 5kx (right)A further magnified imaging of the necking region reveals a network of copper nanoparticles that have sintered at the operating temperature (260 °C) to provide interparticular bonding to the copper powders (Figure 7). It is also clear that the elevated operating temperature did not cause sintering between the larger powders (also witnessed by the loose powder surrounding the printed part following the build). While some nanoparticles are randomly distributed across the copper powder surface, a large amount of copper nanoparticles have concentrated to the necking regions between copper particles.

While some nanoparticles are randomly distributed across the copper powder surface, a large amount of copper nanoparticles have concentrated to the necking regions between copper particles. Nanoparticle bonding was consistently observed in randomly picked regions on both green part surfaces and between layers.Figure 7. Nanoparticle bonding between copper powder particles in copper nanosuspension printed green parts4.3 Green strengthFigure 8 compares the green strength of polymer binder printed parts (cured at 200 °C for1 hour after printing) and copper nanosuspension printed parts (printed with 260 °C bed temperature; directly retrieved from the powder bed without post-printing curing).825614.011.7012.07.80)N10.0( d a o8.0lk a e P6.04.02.00.930.0Polymer binder (100% saturation)Nanosuspension (70% saturation)Nanosuspension (140% saturation)Figure 8.

Green strength of nanosuspension printed parts, in comparison to the polymer binder printed counterparts (evaluated by the rupture loading force in 3-point bending test)The green parts printed by copper nanosuspension with both saturations ratios were ableto be retrieved from the powder bed without breaking. Green parts made with the lower saturation setting (70%) are significantly more fragile than the parts made with 140%. When sufficiently saturated, the nanosuspension printed parts could have a green strength that is comparable to the polymer binder printed parts, which can typically ensure safe handling and depowdering without part damage.4.4 Green and sintered densityThe green density of parts printed with polymer binder and copper nanosuspension iscompared in Figure 9.

Despite the presence of copper nanoparticles at the interstices of the powder bed, the measured green density of parts made via the nanosuspension was not improved substantially.47.0%y t i s i n e d46.0%45.0%45.0%44.8%la n o i t c a r F44.0%43.0%42.0%41.0%Polymer binder (70% saturation)Nanosuspension (140% saturation)Figure 9. Green density of polymer binder and nanosuspension printed partsIn addition to green density, there is also a lack of evidence to show that the sintereddensity is substantially improved by the inkjetted nanoparticles (Figure 10). While increasing powder bed packing through depositing fine particles into powder bed interstices could9257potentially improve sintered density, its effectiveness is limited by the total amount of interstitial particles.

In addition, sintering densification is complicated by many factors including sintering conditions and binder-powder interactions.However, the sintered density difference in 70% and 140% saturated parts suggests that a further increase in binder saturation ratio has a potential to improve sintered density furthermore.

(a) sintered density and (b) sintering shrinkage of polymer binder and nanosuspension printed parts (sintered at 1075 °C for 2 hours)5. Conclusion and future workThis work represents (i) a preliminary investigation into the concept of printing metalsusing the nanoparticles of same material as a permanent and structural binder and (ii) an exploration of the interstitial copper nanoparticle’s effects on green and sintered part properties in Binder Jetting printed copper.With adjusted inter-layer drying and curing conditions (2-minuate interlayer heating at260 °C), the copper nanosuspension can be successfully printed into powder beds without recoating defects and sintering the surrounding powder bed (Section 4.1). The copper green parts printed with copper nanosuspension have obtained satisfactory green strength compared with the polymer binder printed parts without a need for post-print sintering of the powder bed (Figure 8).

Nanoparticle bonding between copper powder particles is evident in nanosuspension printed green parts (Figure 6 and Figure 7).While the copper nanosuspension can be used to successfully print green parts with highquality and strength, the nanoparticle binder’s impact on sintered part properties and sintering10258densification remains uncertain. With the binder saturation ratios (140%) explored in this work, no density increase was observed in green and sintered parts. This ongoing research will focus on expanding the investigated binder saturation ratios, sintering conditions, and base powder packing density to further explore the nanoparticle binder’s impact on sintering.6. AcknowledgementThis material is based upon work supported by the National Science Foundation under Grant No. #1254287.

AcknowledgementThis material is based upon work supported by the National Science Foundation under Grant No. #1254287. Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the views of the National Science Foundation. The authors acknowledge the Binder Jetting technology support provided by ExOne Co, and assistance with 3-point bending test from Mac McCord (Virginia Tech).7. ReferenceAsoro, M., Damiano, J. and Ferreira, P. (2009), “Size Effects on the Melting Temperature ofSilver Nanoparticles: In-Situ TEM Observations”, Microscopy and Microanalysis, Vol. 15 No. S2, p. 706.Bai, J.A., Creehan, K.D. and Kuhn, H.A. (2013), “Powder Particle Layerwise Three-Dimensional Printing Process”.Bai, J.G., Creehan, K.D. and Kuhn, H.A. (2009), “Nanoparticle suspensions for use in the three-dimensional printing process”.Bai, Y., Wagner, G. and Williams, C.B.

nding Mechanisms in Selective Laser Sintering and Selective Laser Melting Prof. Dr. Ir. J.P. Kruth1, Ir. P. Mercelis1, Prof. Dr. Ir. L. Froyen2, Ir. Marleen Rombouts 2 1 University of Leuven (K.U.Leuven), Dept. of Mechanical Engineering, Division PMA 2 University of Leuven (K.U.Leuven), Dept. of Metallurgy and Materials EngineeringReviewed, accepted August 4, 2004AbstractLayer Manufacturing (LM) technologies like Selective Laser Sintering (SLS) were developed in the late 80’s as techniques for Rapid Prototyping (RP). Today, SLS - as well as its derived technology Selective Laser Melting (SLM) - is used as well for prototyping, tooling and manufacturing purposes.

Today, SLS - as well as its derived technology Selective Laser Melting (SLM) - is used as well for prototyping, tooling and manufacturing purposes. This widening of applications is caused mainly by the possibility to process a large variety of materials, resulting in a broad range of physical and mechanical properties.This paper presents a survey of the various binding mechanisms in SLS and SLM, which are responsible for the broad range of materials and applications. Basic binding mechanisms involve solid state sintering, chemically induced binding, liquid phase sintering, partial melting and full melting. Many subcategories can be distinguished based on the type of structural or binder powder composition: single component powder grains (single material or alloy), composite powder grains, mixtures of different powder grains, distinct binder material (sacrificial or permanent), etc.

The paper will explain how these binding mechanisms apply for sintering various types of materials: plastics, metal, ceramics and composites (e.g. glass reinforced polymers, cermets, hardmetals, etc.). It gives a survey of research done at the University of Leuven, Belgium, as well as at other European and non-European organizations.IntroductionSelective Laser Sintering is a Layer Manufacturing process that allows generating complex 3D parts by consolidating successive layers of powder material on top of each other [1,2]. Consolidation is obtained by processing the selected areas using the thermal energy supplied by a focused laser beam. Using a beam deflection system (galvano mirrors), each layer is scanned according to its corresponding cross section as calculated from the CAD model. The deposition of successive powder layers with a typical thickness of 20 till 150 µm is realized using a powder deposition system.Figure 1 shows a schematic example of an SLS system.

Rapid Manufacturing concerns the production of long-term consistent components (e.g. a dental implant) while Rapid Tooling concerns the production of long-term consistent tools (e.g. a plastic injection mould insert).44LaserMirror scannerXY deflectionf- lensθRoll OverflowN Atmosphere2YXFeed containerZ axisBuild cylinderFigure 1A typical SLS machine layoutRM and RT technologies can be classified in a number of ways [3]. Traditionally the technologies are being classified according to the materials they process (plastics, metals, ceramics or composites). In this survey, a classification according to the SLS binding mechanism is used.Figure 2 shows a possible classification of SLS technologies according to the binding mechanism. Most SLS technologies can be classified using this type of classification, however some do not fit in only one section. For example SLS of glass reinforced polymers (e.g.

Most SLS technologies can be classified using this type of classification, however some do not fit in only one section. For example SLS of glass reinforced polymers (e.g. DuraformTM GF being PA12 nylon material with 50 weight % glass) can be seen as Liquid Phase Sintering of composite powder grains regarding the polymer as a binder and the glass particles as a structural material. However, it can also be looked at as if it was a kind of ‘Full Melting’ of a single component composite powder, since the polymer does not just act as a binder material but constitutes the major phase of the structural material being glass filled polymer.

Basically it is a structural material that is ‘fully’ molten, be it that its glass particles remain solid.Next paragraph will shortly describe the different binding mechanisms and illustrate themwith some examples of corresponding technologies.Different binding mechanismsThis paragraph will discuss the state of the art in SLS and SLS-derived technologies. According to the binding mechanism, SLS technologies (and derived technologies like SLM and DMLS) can be classified in four categories. However it has to be noticed that this classification is not absolute and the borders are not always very clear.1. Solid State Sintering (SSS)Solid State Sintering is a thermal process that occurs at temperatures between TMelt/2 and TMelt, where TMelt is the melting temperature of the material. Various physical and chemical reactions occur, the most important being diffusion [4]. It involves neck formation between adjacent powder particles (see Figure 3).

Various physical and chemical reactions occur, the most important being diffusion [4]. It involves neck formation between adjacent powder particles (see Figure 3). The main driving force for sintering is the lowering of the free energy when particles grow together. A gradient in vacancy concentration between the highly curved neck (high vacancy concentration) and the ‘flat’ surfaces (low vacancy concentration) causes a flux of vacancies from the neck and a flux of atoms towards the neck thus increasing the neck size.45Binding mechanism classification1. Solid State Sintering2. Chemically Induced Binding3. Liquid Phase Sintering Partial Melting4.

Solid State Sintering2. Chemically Induced Binding3. Liquid Phase Sintering Partial Melting4. Full Melting3.1 different binder and structural materials3.2 no distinct binder and structural materials4.1 single component single material3.1.1 separate structural and binder grains3.1.2 composite grains3.2.1 single phase material partially molten3.2.2 fusing powder mixture4.2 single component alloyed material4.3 fusing powder mixture3.1.3 coated grainsFigure 2Classification of SLS-technologies according to binding mechanismsThe main advantage of Solid State Sintering is that a wide variety of materials can be processed this way. As long as the temperature is high enough to provide the necessary kinetic energy for the transport of vacancies across the grain boundaries, all powder materials will consolidate via volume diffusion [4].

However, it is a slow process and therefore preheating of the powder material is necessary to increase the diffusion rate of atoms and to obtain an acceptable laser scanning velocity.RDrFigure 3left: neck formation in Solid State Sintering [4], right: neck formation betweentwo stainless steel powder particlesThe National Academy of Science of Belarus, the Baikov Institute of Metallurgy (Russia) and the University of Leuven (Belgium) together investigated Solid State Sintering of titanium powder using ‘soft’ laser radiation [5]. The phenomenon was simulated using numerical models and the results were experimentally justified. To avoid problems related to Selective Laser Melting, titanium powder was consolidated using low intensity laser light (100-300 W/cm2) for a long period of time (5 s). Since the temperature of the particles stayed well below the melting point of Ti, the only mechanism occurring was Solid State Sintering.

Since the temperature of the particles stayed well below the melting point of Ti, the only mechanism occurring was Solid State Sintering. The degree of sintering was quantified using the ratio of neck diameter to powder particle diameter (D/2R in Figure 3). It was concluded that the degree of sintering was low as long as the temperature reached stayed below the α/β-transition temperature (890°C). Rising above this temperature, thus forming β-phase titanium resulted in a strong rise in sintering degree. This could be explained by the high self- diffusion coefficient of the β−phase compared to the α−phase.46Phenix Systems (France) commercializes an SLS-like system that realizes Solid State Sintering using a high temperature process chamber reaching up to 900 ˚C [6]. The compacted powder, which is close to the Solid State Sintering temperature, is sintered due to the energy contribution of a Nd:YAG laser source. This way, ceramic materials can be processed as well as metal powders.

This way, ceramic materials can be processed as well as metal powders. To obtain the desired characteristics, a post sintering operation is necessary. Figure 4 depicts some ceramic parts.Figure 4Solid state sintered ceramic parts (source: Phenix Systems)2. Chemically Induced BindingAt the Fraunhofer Institute IPT (Aachen, Germany) the direct production of SiC ceramic parts has been investigated [7]. No binder elements were used and the laser-material interaction times were very short, thus excluding the diffusion processes occurring in Solid State Sintering. When heating the SiC particles to a very high temperature, partial disintegration of the SiC into Si and C occurs. The free Si forms SiO2, which acts as a binder between the SiC particles. The parts are thus composed of a mixture of SiC and SiO2. Afterwards an infiltration step using Si yields full dense parts.Next to SiC, also other ceramics were tested, e.g. Al2O3 and ZiSiO4.

The parts are thus composed of a mixture of SiC and SiO2. Afterwards an infiltration step using Si yields full dense parts.Next to SiC, also other ceramics were tested, e.g. Al2O3 and ZiSiO4. Figure 5 shows a ceramiccasting shell produced by SLS.An other example of chemical induced binding is the consolidation of Al powder in a N2 atmosphere whereby an AlN binder phase is formed. This reaction is not yet used in Selective Laser processing, but Sercombe used it in an SLS post-curing process, after polymer debinding and before infiltration [15] (see paragraph 3.1.3 Coated grains).Figure 5SLS-produced investment casting shell for impeller wheel (source: FraunhoferInstitute IPT)473. Liquid Phase Sintering (LPS) – Partial MeltingThis category unites many different kinds of technologies. Most of these technologies combine a structural material remaining solid throughout the process and a binder material being liquefied.

Most of these technologies combine a structural material remaining solid throughout the process and a binder material being liquefied. In some cases, however, the solid and the liquid phases result from the same material. A first group of technologies is characterized by a clear distinction between the binder and structural materials.3.1 Different binder and structural materialsTechnologies can be further divided into three groups according to the type of powder grainsthat is used.3.1.1 Separate grainsThese technologies use different binder and structural grains. The structural material can be a metal as well as a ceramic material; the binder materials however are mostly metals. Binder particles are usually much smaller then the structural ones.

The structural material can be a metal as well as a ceramic material; the binder materials however are mostly metals. Binder particles are usually much smaller then the structural ones. Therefore, the high surface-to-volume ratio of the grains – in combination with its relatively low melting point – favors the preferential melting of the binder material [8].The combination of small binder particles and larger structural particles has the additional benefit of better packing with small pores favoring the spreading of the liquid binder by capillary forces and causes a rearrangement of the particles. Generally a ‘green’ part is produced which is still porous and brittle. Therefore, a post treatment consisting of a furnace post-sintering, Hot Isostatic Pressing (HIP) or an infiltration with a low melting point material is usually necessary. The last solution is most commonly used because of its effectiveness and its relative ease.Figure 6 shows the LPS mechanism schematically.

The last solution is most commonly used because of its effectiveness and its relative ease.Figure 6 shows the LPS mechanism schematically. The solid line represents a conventional furnace sintering process. The LPS mechanism, as it occurs in SLS, is represented by the first part of the dashed line, since only the rearrangement phase takes place. The process is frozen at this stage resulting in a porous green product. A thermal post-treatment in a furnace may be used to complete the cycle.Figure 6Mechanism of Liquid Phase SinteringMany different material combinations have been tested in the past. At the University of Leuven, metal-metal composites were tested as well as metal-ceramic composites [9,10]. Some examples of metal-metal composites are Fe-Cu and Stainless Steel-Cu. As for the metal-ceramics combinations, WC-Cu, WC-Co, WC-CuFeCo, TiC-Ni/Co/Mo, ZrB2-Cu and TiB2-Ni were tested.

Some examples of metal-metal composites are Fe-Cu and Stainless Steel-Cu. As for the metal-ceramics combinations, WC-Cu, WC-Co, WC-CuFeCo, TiC-Ni/Co/Mo, ZrB2-Cu and TiB2-Ni were tested. Figure 7 and Figure 8 left show cross sections of a Stainless Steel-Cu and a WC-Co green part.48cba100 mµFigure 7LPS of Stainless Steel-Cu powder mixture (a: non molten steel particle, b: moltenCu, c: porosity) (source: University of Leuven)100 mµLPS of WC-Co powder mixture, left: before infiltration (a: non molten WC Figure 8 particle, b: molten Co, c: porosity), right: after infiltration with copper (source: University of Leuven)3.1.2 Composite grainsComposite powder grains contain both the binder and the structural material within each individual powder grain.

The right side picture shows a low-weight, high-strength sandwich panel, which was made from the glass filled nylon material.50Figure 11Left: cross section of a glass filled nylon part, right: light-weight sandwich panelmade from glass filled nylon (source: University of Leuven)3.1.3 Coated grainsA third possibility to combine a binder and a structural material is to coat the structural material with the binder phase. This ensures that the laser radiation hitting the powder particles is preferentially absorbed by the binder material that is to be melted: in the case of separate grains it may happen that the structural material (e.g. Fe) absorbs more laser energy than the binder material (e.g. Cu) causing the structural material to melt before the binder does [12]. Moreover a more effective bonding of the structural particles is realized since the binder material already surrounds all structural particles.

Moreover a more effective bonding of the structural particles is realized since the binder material already surrounds all structural particles. Coated powders exist both with metal and polymer binder coatings.At the University of Leuven, experiments with Cu coated steel powder turned out to besuccessful despite of the fact that the powder coating process was hard to control [4].Polymer coated steel powders (DirectSteel and LaserForm) are commercially produced by 3D Systems [13], formerly DTM Corporation. Most polymer coated powders need an additional debinding step in which the polymer is burnt out. Afterwards an infiltration step is possible. Figure 12 shows a cross section of an SLS part made from polymer coated stainless steel powder (LaserForm), after polymer debinding and infiltration with bronze.Figure 12Bronze infiltrated Laserform ST 100 partHowever, polymer coated powders may also be applied without debinding in order to produce e.g. polymer/metal composite parts.

polymer/metal composite parts. An example of this is polyamide coated copper51powder used on DTM machines to produce copper reinforced polyamide moulds for injection moulding [14].At the University of Queensland, an aluminium sintering process was developed in collaborations with 3D Systems [15, 16]. Since the strong oxide layer on the grains prohibits direct sintering of Al powder, an indirect procedure was developed. Nylon coated Al grains are fused together using the LPS mechanism. Next, the nylon binder is decomposed. A small amount of magnesium added to the powder mixture makes it possible to form aluminium nitride instead of aluminium oxide when heating up to 540 ºC. The aluminium nitride acts as a rigid skeleton, which is necessary during the infiltration phase. The infiltrant being used is a ternary eutectic aluminium alloy (Al-13.8Si-4.7Mg), having a lower melting point than the basic aluminum alloy. Relative part densities up to 95 % were obtained.

The infiltrant being used is a ternary eutectic aluminium alloy (Al-13.8Si-4.7Mg), having a lower melting point than the basic aluminum alloy. Relative part densities up to 95 % were obtained. Figure 13 shows the different processing steps.a50 mµb50 mµc20 mµFigure 13Steps in Aluminium SLS: a) green part (nylon binder not visible) b) aluminiumnitride skeleton surrounding the aluminium grains, c) infiltrated partA last type of coated powders are the sand powders developed DTM- 3D Systems (SandForm) as well as by EOS GmbH (Direct Croning) [17]. The powders offered are standard Si of Zr sands, coated with a small phenolic layer, responsible for fusing the sand grains together. These powders are used to produce moulds and cores for metal casting.3.2 No distinct binder and structural materialTechnologies in this group do not exhibit a clear distinction between binder and structural phases.

Rather than the distinction between binder and structural material, there is a distinction between molten and non-molten material areas. Therefore ‘Partial Melting’ is a better name for these technologies than ‘Liquid Phase Sintering’.3.2.1 Single phase, partially moltenWhen the heat supplied to a powder particle is insufficient to melt the whole particle, only a shell at the grain border is molten. The core of the grain stays unaffected. This way the molten material acts as a binder between the non-molten particle cores. This binding mechanism can arise as well with metals as with polymers, although the consolidation of polymer powders results also from other mechanisms (consolidation at the glass transition temperature, which is lower than the melting temperature, polymer chain rearrangement and cross-linking).The partial melting phenomenon was modeled by Karapatis [8]. Using a simple thermal model, skin and core temperatures of the powder particles were calculated.

Using a simple thermal model, skin and core temperatures of the powder particles were calculated. This way the minimal52pulse energy to fully melt the particle can be calculated. Below this value, the core temperature never exceeds the melting temperature and only partial melting is obtained.3.2.2 Fusing powder mixturePowders consisting of multiple phases can be classified in this group when they are only partially molten. For example at the University of Leuven experiments were done with a Fe- Fe3P-Ni-Cu powder mixture, aiming at the production of full density parts [18]. The addition of a melting point lowering additive like Fe3P is favourable in making the process more energy efficient. For instance, alloying pure Fe with a small amount of P lowers the melting point of pure Fe (1538°C) to the eutectic temperature of the Fe-P system (1048°C, see Figure 14 ).

For instance, alloying pure Fe with a small amount of P lowers the melting point of pure Fe (1538°C) to the eutectic temperature of the Fe-P system (1048°C, see Figure 14 ). Moreover, the dissolution of P in Fe also has the benefit of a lower surface tension of the melt, resulting in a better wetting behaviour. Ni is added for its marked strengthening effect.Figure 14Phase diagrams of Fe-P and Cu-P showing the temperature lowering effect of theadded PA closer look at the microstructure reveals that not all the powder particles are molten (see Figure 15 ). The final part consists of a low melting point P-rich phase (no. 3), a high melting point phase with no significant amount of P (no. 2), some remaining porosities (no. 4) and some remaining unmolten Fe powder particles (no. 1).

3), a high melting point phase with no significant amount of P (no. 2), some remaining porosities (no. 4) and some remaining unmolten Fe powder particles (no. 1). Therefore, the process cannot be called ‘full melting’ and the name ‘partial melting’ is preferred.Figure 15 Micrograph of multiphase steel powder: 1) unmolten Fe particle, 2) high melting P-poor phase, 3) low melting P-rich phase, 4) pores (source: University of Leuven)53EOS’ Direct Metal Laser Sintering process also belongs to this category [17]. The recent introduction of 20 µm metal powders brings an improvement in mechanical properties and surface quality. Figure 16 and Figure 17 show the difference in porosity between a 50 µm and a 20 µm powder. In both cases shot penning enhances surface quality and improves the density of the top layer.

Figure 16 and Figure 17 show the difference in porosity between a 50 µm and a 20 µm powder. In both cases shot penning enhances surface quality and improves the density of the top layer. Different parameter sets are used for the part’s skin and core, in order to obtain a rigid and hard shell and at the same time a higher building speed (by fast scanning of the core).ab200 mµcd100 mµFigure 16 DMLS Direct-Metal 50V-2 grain size 50 µm, porosity 10-15%: a and c before shot-penning, b and d after shot peening (a and b: cross section, c and d: top view)(Source EOS)1 mmFigure 17 DMLS Direct-Steel 20V-2 grain size 20 µm, porosity ~ 5%, in the ZX plane: left before and right after shot-penning, cross section (Source IPT)Table 1 concludes the third bonding category by summarizing some important mechanicalproperties of commercially available materials.54ReferenceSLS materials0 2 P : l e e t S e c n e r e f e R5 . 0 u C g M n Z 5 7 0 7 lA e c n e r e f e R0.

Full MeltingDriven by the need to produce near full dense objects, with mechanical properties comparable to those of bulk materials and by the desire to avoid lengthy post processing cycles, Selective Laser Melting has been developed. Polymers as well as metals can be completely molten by a laser beam. However, the appellation Selective Laser Melting (SLM) is reserved for metallic materials. Therefore, polymers will not be discussed in this section.All metals may be thought as candidate material however large differences exist in the ease of processing (e.g. laser absorption, surface tension and viscosity of the liquid metal, etc.). For each new material, a process-window needs to be determined experimentally, in order to ovoid scan track instabilities (sphereodisation of the liquid melt pool, also known as ‘balling’) and part porosity. Therefore the range of commercially available SLM metals is still limited today.

Therefore the range of commercially available SLM metals is still limited today. Figure 18 shows an example of such a process window for an iron based powder mixture.55Figure 18SLM process window for an iron based powder mixture (source: University ofLeuven)4.1 Single component, single material powderThe Fraunhofer Institute ILT [19] as well as the Universities of Osaka [20] and Leuven, performed a lot of investigation on Selective Laser Melting of pure titanium powder. Figure 19 shows a titanium part that was produced with SLM. It can be seen that the part density is almost 100%, however the picture also reveals a crack (bottom left) caused by thermal stress, which needs to be reduced in the future.

It can be seen that the part density is almost 100%, however the picture also reveals a crack (bottom left) caused by thermal stress, which needs to be reduced in the future. It is expected that commercial titanium SLM will be offered soon.Figure 19Cross section of a Ti part produced by Selective Laser Melting (source: Universityof Leuven)4.2 Single component, alloyed powder particlesThese powders consist of a single kind of grains, consisting of alloyed materials. Many different kinds of alloys have been tested and some are already commercially available trough the three German manufacturers of SLM machines (Concept Laser GmbH [21], Trumpf [22] and MCP-HEK [23]). The Fraunhofer Institute ILT for example has investigated Stainless Steel 1.4404 and Tool Steel 1.2343 [24].

The Fraunhofer Institute ILT for example has investigated Stainless Steel 1.4404 and Tool Steel 1.2343 [24]. Figure 20 shows an example of a stainless steel part.Next to the steel powders, also a titanium alloy (TiAl6V4) was tested for the manufacturingof medical implants.56Figure 20Cross section of a stainless steel 1.4404 part produced by Selective Laser Melting(source: Trumpf)Figure 21 compares the mechanical properties of parts made of steel 1.4404 with those of the bulk material. It can be concluded that mechanical properties are comparable to bulk material apart from the ductility, which is strongly reduced. This might be overcome applying the right heat treatment.

The powders and sintered parts depicted in Figure 15 illustrate that the border between ‘partial melting’ and ‘full melting’ is rather vague, at least when near-full density parts are aimed for.ConclusionsOver the last decade Rapid Prototyping techniques gained a wide acceptance. Among the different RP techniques, Selective Laser Sintering had the advantage of being able to process a wide range of materials. This large variety of materials gradually allowed to extend the field of applications, from simple visual prototyping to functional prototyping, and even Rapid Tooling and Rapid Manufacturing applications.In recent years, the idea of full melting metal powders was explored, supported by the continuously improving process parameters (smaller layer thickness, smaller spot size, etc.), resulting in mechanical properties being much better than those of early time selective laser sintered parts and comparable with bulk metal properties.

Rapid Tooling and specially Rapid Manufacturing technologies are still limited in use, but they are steadily growing. It can be expected that in the next decade they will achieve the same level of acceptance as Rapid Prototyping, by continuous improvement of material properties and accuracy and by a decrease of the costs.References[1] J.P. Kruth, B, Van Der Schueren, J.E. Bonse, B. Morren, Basic powder Metallurgical Aspects in Selective Metal Powder Sintering, Annals of the CIRP Vol. 45/1, 1996 [2] J.P. Kruth, X. Wang, T. Laoui, L. Froyen, Lasers and Materials in Selective Laser Sintering, Proc. Int. Conf. on Laser Assisted Net Shape Engineering, 2001 [3] G.N. Levy, R. Schindel, J.P. Kruth, Rapid Manufacturing and Raid Tooling with Layer Manufacturing (LM) Technologies, State of the Art and Future Perspectives, Annals of the CIRP, Vol. 52/2, 2003 [4] B.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperBIO-INSPIRED AM STRUCTURE WITH CARBON FIBRE REINFORCEMENTP. Jameekornkul*, A. Panesar*IDEA Lab, Department of Aeronautics, Imperial College London, UKAbstractThis study presents a novel design approach that introduces the bouligand fibre arrangement in a honeycomb structure. This design is made possible by utilising AM technology which allows for the realisation of complex design whilst ensuring precise control for fibre placement. Reinforced honeycomb structures with varied incrementally twisted pitch angle of 0°,1.25°,5° and 15° were successfully fabricated and tested under out- of-plane compression. Among the tested angles, the honeycomb structure oriented at 1.25° exhibited higher peak load and higher specific energy absorption.

Among the tested angles, the honeycomb structure oriented at 1.25° exhibited higher peak load and higher specific energy absorption. Unlike the original untwisted honeycomb, the bio-inspired honeycomb showed no noticeable buckling or delamination at the mid-plane, which may be due the microcracking or resulting from a delay in crack propagation. However, the precise relationship between pitch angle and honeycomb properties requires further investigation across a wider range of pitch angle with more focus understanding the fracture propagation in fibre arrangement. Overall, the preliminary results indicate that the proposed bio-inspired AM design present a promising approach to enhance the properties of honeycombs and afford the flexibility to improve toughness and energy absorption capabilities.IntroductionBouligand structureOver the course of many years, nature has evolved structures that possess remarkable characteristics.

Bio- inspired structures have gained significant attention in recent years due to their remarkable energy absorption capabilities and their potential applications in areas such as damage mitigation. One of bio-inspired structure which was adapted in manufacturing the composite structure is Bouligand structure introduced by Bouligand in 1972 [1] or the helicoidal formation. The formation consisting of unidirectional fibrous layer orientating around the centre with a constant pitch angle (γ) throughout the stack of layers [2]. This type of structure arrangement could be found in natural microstructure and have been studied by researchers, for example, mantis shrimp [3], DNA structure, the cuticle of scorpion chela [4], palm leaf [5] [6], arthropod phylum [7] and bone minerals [7].Figure 1: The schematic of helicoidal fibre structure where γ is a constant pitch angle.

[8]The helicoidal fibre structure has been shown to provide high interfacial strength, maximise transverse load resistance [13-14], delaying the fracture propagation, higher damage resistance compare to conventional cross ply [11], energy absorption ability, remarkable toughness and the impact strength [4] [12] [13]. These1738features associated with the twisting of reinforcement fibre along with the stacking direction and therefore, force the crack to twist as they grow parallel to the fibre orientation [14]. Another feature of helicoidal structure is creating microcracks within the structure and increase in the surface crack area which responsible for more energy able to be absorbed before the catastrophic failure [14]. Apart from experimental investigation, theoretical models were proposed for optimal pitch angle [15] and crack growth behaviour [14]. So far, most helicoidal structure studies focused on a laminate lay-up which limited the geometric freedom.

So far, most helicoidal structure studies focused on a laminate lay-up which limited the geometric freedom. Therefore, this study aims to utilise the capabilities of AM technique to further control the helicoidal inspired fibre structure.Failure mechanism in fibre compositeFigure 2: Major failure modes observed for the fibre reinforced composite [16]Major failure mechanisms occurred within composite with fibre reinforcement are delamination, propagation of the matrix cracking and fibre failure [17], [18] (Figure 2).

It can be summarised from a review [16] of composite structure that different failure modes were observed in varied mode of fracture classification (Mode I, II and III) and different composite shapes e.g., tubes, sections, and cellular structures resulting in changing of the crashworthiness performance.While AM technology has not been commonly used and studied for composite thin-wall tube, the filament winding and lay-up technique have known for mass production of carbon fibre-reinforced polymer (CFRP) tubes which normally have reinforcement layers at the outer skin. Ma et al. [18] conducted an axial crushing on a hybrid composite tubes suggested that the CFRP circular tube absorbed energy through fibre fracture, matrix failure and interlayer fracture [19] which demonstrated through the outer layer of reinforcement rolled out while the inner layer is bent and fractured [18] [20].

The study confirmed the failure mechanism and progressive collapse behaviour are the key to crash efficiently of thin-walled composite tube. Thus, here, we will examine the failure mechanisms that occur in our modified structure to understand crashworthiness performance at the root causes.Figure 3: a.) Interlaminar shear stress distribution predicted in [15] varying the pitch angle b.) the optimal rotation angle for different total number of laminae [15]1739AimsHoneycomb characterised by their hexagonal cell structure which capable of optimising the space filling, are widely used in sandwich structures to provide superior strength and stiffness while maintaining a lightweight profile. In this study, our design involves a modified honeycomb structure with a helicoidal fibre arrangement. The objective is to leverage the advantages offered by bio-inspired structures, specifically focusing on enhancing energy absorption capability and improving fracture toughness.

The objective is to leverage the advantages offered by bio-inspired structures, specifically focusing on enhancing energy absorption capability and improving fracture toughness. Our aim is to explore novel approaches for achieving controlled crack propagation by utilising the continuous additive manufacturing (AM) technique, which enables complex part design and tailorable fibre placement.Therefore, we will investigate the compression properties of this bio-inspired honeycomb structures, focusing on their potential for energy absorption applications. experiments will be conducted to assess the feasibility and effectiveness of the proposed design concept. Specifically, the mechanical performance and energy dissipation capabilities of a thin-walled tube with a helicoidal-honeycomb profile will be evaluated.

Specifically, the mechanical performance and energy dissipation capabilities of a thin-walled tube with a helicoidal-honeycomb profile will be evaluated. These experiments will provide insights into the performance characteristics of the bio-inspired structure and its potential for extending the concept to further applications.MethodologyDesign processA single honeycomb structure was designed using the Solidworks software, and its dimensions are presented in Figure 4a. The structure consists of 48 layers and has a height of 17.7 mm. Figure 4a demonstrates the twisting of the honeycomb by shifting the XY coordinates based on the pitch angle θ around the fixed centre point (a, b). The pitch angle was varied at intervals of 1.25°, 5°, and 15°, resulting in 1, 4, and 12 full rotations within the sample, respectively.

The pitch angle was varied at intervals of 1.25°, 5°, and 15°, resulting in 1, 4, and 12 full rotations within the sample, respectively. Figure 4b is an example of travel path of the printer nozzle when the honeycomb was shifted using a 15° pitch angle.a.) b.)Figure 4: a.) Dimension of honeycomb (top view) b.) trajectory of nozzle for sample with 48 layers and 15 degrees pitch angleManufacturing processThe material extrusion technique is chosen for the fabricating of the samples using Anisoprint, a printer with capability of co-extruding the fibre reinforcement and matrix. The G-CODE, responsible for controlling the nozzle's path and the amount of material extrusion, was modified to achieve the desired helicoidal fibre arrangement in the printed structure.1740To ensure the structural integrity, the limitation of printing process was determined. The minimum length of the inner side of the honeycomb should at least 7.5 mm to avoid the fibre curvature and fibre twisting.

The minimum length of the inner side of the honeycomb should at least 7.5 mm to avoid the fibre curvature and fibre twisting. In addition, the preliminary printing test was conducted to determine the appropriate volume fraction between the carbon fibre reinforcement and matrix for the printing. The volume fraction plays a significant role in the properties of the printed specimen, as the quality of bonding between the matrix and the fibres directly depends on this parameter.During the manufacture, the model dimensions, volume fraction and the amount of material extruded were kept unchanged. The matrix material used was PETG, while carbon fibre was employed as the continuous fibre reinforcement (refer to Table 1 for combined material properties). To maintain continuity, the fibre reinforcement was only cut once at the end of each layer. The feed rate and travel speed were set to the default values provided by the Aura slicer software.

To maintain continuity, the fibre reinforcement was only cut once at the end of each layer. The feed rate and travel speed were set to the default values provided by the Aura slicer software. The nozzle temperature and build plate temperature were consistently maintained at 245°C and 60°C, respectively.Table 1: Carbon fibre and matrix properties [21]Materials CCF-1.5K-PETGCompressive Strength (MPa) Compressive Modulus (GPa)237.4 ± 4.249.0±2.4Figure 5: Printed specimen (15°, 5°, 1.25° pitch angle from right to left)Figure 5 displays the printed specimens with varying pitch angles. It is important to mention that the default slicer settings resulted in an excess of PETG material at the start of each layer, which could be removed through post-processing. However, to maintain the integrity of the bonding properties, post-processing was not performed.

However, to maintain the integrity of the bonding properties, post-processing was not performed. To ensure fairness in the experiment, the weight of each specimen was measured, and it was confirmed that an equal amount of excess plastic was present at each layer in all specimens and will not impact the experimental results.ExperimentThe out-of-plane compression test was carried out using the 50kN Instron 5969 machine, with a constant displacement rate of 1mm/min. The video camera was used to record the experiment in order to observe the failure mechanism along with time mark. For each sample, three specimens were tested. Additionally, a video camera was used to record the test, facilitating fracture investigation. Following the compression test, the failed specimens will be examined using a scanning electron microscope (SEM), with a focus on observing crack initiation and propagation.

Following the compression test, the failed specimens will be examined using a scanning electron microscope (SEM), with a focus on observing crack initiation and propagation. It is important to note that the experiment was conducted within 72 hours after the fabrication process to minimize the potential impact of moisture absorption in PETG.In order to further the understanding of compressive behaviour, the specific energy absorption (SEA)will be calculated from Equation (1).𝑆𝐸𝐴 =𝐸𝑚(Equation 1)1741where 𝐸 is the total energy absorbed which equal to area under force-displacement curve following equation 2 and 𝑚 is mass of material (kg) which is an average weight of all printed specimens (0.39 g).𝑑𝑚𝑎𝑥 𝐸 = ∫ 0𝐹𝑑𝑠(Equation 2)where 𝑑𝑚𝑎𝑥 is the maximum displacement achievable by all specimens. To distinguish between the elastic and plastic phase, the yield strength determined by the stress value which corresponding to the 0.2% offset strain.

To distinguish between the elastic and plastic phase, the yield strength determined by the stress value which corresponding to the 0.2% offset strain. The highest stress value of the stress-strain curve before reaching the densification phase at the end will be defined as the ultimate strength.Results and DiscussionsStress-strain curve and pitch angle effectIn Figure 6, the stress-strain plot demonstrates a similar pattern of ductile materials behaviour across the range of pitch angle. The specific energy absorbed, ultimate strength and yield strength are calculated and illustrated in according to pitch angle in Figure 7. The experiment showed that the orientation of reinforcement in each layer affects both peak force and specific energy absorbed (Figure 7). The rotation angle of 1.25° has shown the highest energy absorbed by achieving the highest ultimate strength and the smallest stress drop level, indicating a less catastrophic shear failure at the end.

While the rotation angle of 15° showed higher peak load and modulus compared to the normal honeycomb but experienced a significant drop in stress afterwards.Figure 6: Stress-Strain plot for compression experiment varying pitch angle.Figure 7: Properties of varying pitch angle calculated from displacement-load curve.1742Both specimen of 5° and 15° orientations reached peak strength at lower strain than the normal honeycomb and leads to a lower stress before entering densification phase.

Although no clear trend was observed between properties and the pitch angle, our most successful samples, specifically those with a 1.25° rotation angle, align with the natural bioligand structure observed at approximately 1.6° and supported by studies [22] [23] which indicate that indicate superior performance with smaller pitch angles under their conditions.The potential cause of performance fluctuation at different pitch angle is suspected to be attributed to the interlaminar shear behaviour. This interlaminar shear distribution was affected by the fibre stacking orientation, ply thickness and total number of stacking ply [15]. The prediction model proposed by [15] showed that plateau stress was not linearly proportional to pitch angle (Figure 3b and c).

The prediction model proposed by [15] showed that plateau stress was not linearly proportional to pitch angle (Figure 3b and c). In addition, [23] suspected that the fluctuation could be affected by the fibre intersection points changed by pitch angle causing the stress and fracture to propagate differently.It is important to note that this study introduces the honeycomb structure alongside with AM technique, which potentially contributes to the observed fluctuations due to manufacturing and complex design effect causing inconsistence fibre spacing and layer thickness [14]. In addition, the number of full rotations back to original first layer should be taken in account.

In addition, the number of full rotations back to original first layer should be taken in account. Therefore, a broader range of pitch angle and an effect of rotation rounds was planned to be investigated further.Failure modes0°1.25°5°15°Figure 8: The failure behaviour during the quasi-static compression test captured by a video camera (Pitch angle 0°, 1.25°, 5°, 15° from top to bottom).1743Figure 8 and Figure 9a show the sign of buckling and delamination in the mid-plane of the print direction of the zero-degree honeycomb. This observation aligns with the findings of Zhang and Ashby [24] which expected the honeycomb to buckle elastically. However, the samples with positive pitch angle did not clearly show the delamination or buckling, this could be due to the microcracking occurring inside of the structure, or the progressive collapse at slower rate, or the buckling was delayed and cannot be captured by video camera.

Further in-situ measurement where the load will be stopped for observation could be conducted.Apart from general trend of stress-strain curve, all specimens shown similar shear failure (see last column of Figure 8) starting from the mid-plane progressing towards the bottom layer (Figure 9b and c) after the upper yield stress was reached. The shear failure can be clearly seen in SEM images that both fibre and matrix shown catastrophic failure in the said direction. This shear catastrophic failure then causes the delamination between interfaces before entering the densification phase at the end.a.)b.)c.)Figure 9: Failed specimens (first row) and SEM images (second row) show a.) buckling in mid-plane of original honeycomb b-c.) Shear failure at the bottom layer.ConclusionThe preliminary results obtained from the compression test revealed a consistent trend in the stress-strain curve.

However, the ultimate strength and the point at which stress dropped due to shear failure varied with different pitch angles. The orientation and alignment of the fibres were influenced by the pitch angle, leading to potential differences in crack propagation rates and the presence of invisible microcracks. Consequently, the mechanical properties exhibited fluctuations depending on the pitch angle, requiring further investigation to establish a prediction model. A broader range of pitch angles should be explored in future studies, and the compression test could be paused at different phases to gain a deeper understanding of fracture mechanics.In terms of quality inspection, shear failure at both fibre and matrix was observed in all specimens, while only the conventional honeycomb showed signs of buckling and crack propagation.

O-INSPIRED DESIGN OF BIPOLAR PLATE FLOW FIELDS FOR POLYMER ELECTROLYTE MEMBRANE FUEL CELLSNannan Guo1, Ming C. Leu1, and Maoliang Wu21Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409, USA 2Shanghai University of Electric Power, Shanghai, 200090, ChinaAbstractThe flow field of a bipolar plate distributes hydrogen and oxygen for polymer electrolyte membrane (PEM) fuel cells and removes the produced water from the fuel cells. It greatly influences the performance of fuel cells, especially regarding reduction of mass transport loss. Flow fields with good gas distribution and water removal capabilities reduce the mass transport loss, thus allowing higher power density.

Flow fields with good gas distribution and water removal capabilities reduce the mass transport loss, thus allowing higher power density. Inspired by natural structures such as veins in tree leaves and blood vessels in lungs, which efficiently feed nutrition from one central source to large areas and are capable of removing undesirable by-products, a mathematic model has been developed to optimize the flow field with minimal pressure drop, lowest energy dissipation, and uniform gas distribution. The model can be used to perform optimal flow field designs, leading to better fuel cell performance for different sizes and shapes of bipolar plates. Finite element modeling (FEM) based simulations and in-situ experiments were conducted to verify some of the flow field designs obtained using the developed mathematic model.1.

Finite element modeling (FEM) based simulations and in-situ experiments were conducted to verify some of the flow field designs obtained using the developed mathematic model.1. IntroductionPolymer electrolyte membrane (PEM) fuel cells have great advantages as an energy conversion device such as low-temperature operation, high-power density, fast start-up, system robustness, and low emissions. They are very promising alternative power sources for automotive applications [1]. The flow field of a bipolar plate supplies fuel and oxygen to reaction sites and removes reaction products (i.e., liquid water) out of the fuel cell, significantly effecting the performance of PEM fuel cells. Previous studies have demonstrated that the power density of a fuel cell can increase by 20-30% with a proper flow field design [2, 3].Numerous flow field designs have been proposed and investigated.

These designs can be classified into four categories of layouts: pin-type, parallel, serpentine and interdigitated [4]. Among them parallel and serpentine designs are the most widely known and utilized. Parallel designs usually provide low pressure drop, and serpentine designs usually give high fuel cell performance. However, lots of drawbacks are associated with these two kinds of designs. Parallel flow fields have multiple flow paths. Because gases flow preferentially through the least resistant channels, stagnant areas tend to form at various areas of parallel channels, lowering the transport efficiency of the reactants. Serpentine flow fields [5] typically result in a relatively long reactant flow path, leading to a substantial pressure drop.

Serpentine flow fields [5] typically result in a relatively long reactant flow path, leading to a substantial pressure drop. Consequently, significant parasitic power loss is associated with the fuel supply (as high as over 35% of the stack’s output power), especially when the stack contains a large number of cells or when fuel cells with large active areas areused. Additionally, reactant concentration significantly decreases from the inlet to the outlet of the flow channel. This decrease leads to considerable Nernst losses and non-uniform current density distribution, reducing both the performance and the lifetime of PEM fuel cells [6].Besides conventional flow field layouts, researchers have begun to investigate flow field designs with inspirations from nature [7-12]. These bio-inspired flow field designs have demonstrated advantages in distributing reactants and enhancing performance. Kloess et al.

These bio-inspired flow field designs have demonstrated advantages in distributing reactants and enhancing performance. Kloess et al. [7] combined serpentine and interdigitated patterns with a leaf/lung layout to design a bio-inspired flow layout. They reported that these new designs improved pressure distribution, decreased pressure loss, and increased power density as compared to pure serpentine and interdigitated designs. Chapman et al. [8, 9] claimed that their bio-mimetic designs enhanced the performance of PEM fuel cells by 16%. They stated that their designs comprised main and sub-feed channels and that tapered, non-linear channels could improve performance. The above studies have shown the potential benefits of incorporating natural structures into flow field designs. However, the bio-inspired designs that have been investigated thus far are only bio-mimetic, i.e. they mimic some particular structures in nature’s biological systems.

However, the bio-inspired designs that have been investigated thus far are only bio-mimetic, i.e. they mimic some particular structures in nature’s biological systems. There is a lack of comprehensive and systematic studies on how to create an optimal bio-inspired design of flow fields for bipolar plates.Very recently, some researchers have developed optimization models to optimize flow field designs in order to improve the performance of PEM fuel cells. Wang et al. [13] integrated a conjugated-gradient scheme and a three-dimensional fuel cell numerical model to optimize the depth of tapered flow channels for serpentine designs, which enhances the oxygen transport rate and local current density. Peng et al. [14] proposed an analytical model to optimize the depth of flow channels in a slotted-interdigitated flow field in order to obtain an even flow distribution in the channels.

Peng et al. [14] proposed an analytical model to optimize the depth of flow channels in a slotted-interdigitated flow field in order to obtain an even flow distribution in the channels. However, all these optimization models were developed based on a certain flow layout (serpentine or interdigitated). Moreover, only the channel’s width or depth is optimized to achieve the expected flow properties, which limits the optimization space of flow field designs.With inspiration from the idea that nature structures (e.g., leaf veins) provide excellent mass transport with uniform pressure distribution and low energy dissipation, the present study presents an optimization model to generate flow field designs with these properties. The generated designs include designs of flow field layouts and channel geometries. The model starts with small grids on the flow fields and generates optimum designs with structures similar to nature.

The model starts with small grids on the flow fields and generates optimum designs with structures similar to nature. This is intended to lead to minimum pressure drop, uniform pressure distribution, and better fuel cell performance. Given the flow field shapes and the locations of the inlet and outlet, one example of the generated designs based on the optimization model has been investigated with FEM simulation and compared with the conventional parallel and serpentine designs. The results show that the proposed design provides better fuel cell performance, with more uniform pressure distribution and less pressure drop over the flow field.2. Mathematical Algorithm 2.1 Flow fieldThe flow field design involves the layout of the flow field and channel cross-section geometries including shape and dimensions.

Mathematical Algorithm 2.1 Flow fieldThe flow field design involves the layout of the flow field and channel cross-section geometries including shape and dimensions. Usually for the flow field design, researchers startfrom a particular flow field layout (e.g., serpentine [13] or interdigitated [14]) and then optimize the shape or dimensions (e.g., width and depth) to achieve desired results. Our approach starts from the flow field grids (see Fig. 1), where the flow field is meshed by × nodes channels with pre-set dimensions. After optimizing the channel and dimensions, the corresponding flow field layout is determined because the flow field layout essentially is the configuration of all the flow channels.× × + ×2.2 Optimization modelThe mathematic model is formatted to a constrained optimization problem.

The steps for solving the problem are as follows: (1) Mesh the flow field into small grids, which can be triangle, rectangular, or hexagonal; (2) Establish flow relationships using these grids, and define the objective function and constraints; (3) Optimize the objective function using the Lagrangian method; and (4) Generate the optimum flow field design. The flow chart of the optimization model is given in Fig. 1.Fig. 1 Flow chart of the optimization model.2.2.1 Flow relationshipIn the proposed optimization model, several assumptions are made. Firstly, the flow through the channels is laminar flow. This assumption is fairly true in reality, since, under most conditions, gas flow in fuel cell channels is laminar; only at extremely high flow rates can gas flow become turbulent [15].

This assumption is fairly true in reality, since, under most conditions, gas flow in fuel cell channels is laminar; only at extremely high flow rates can gas flow become turbulent [15]. Secondly, only the pressure drop along each channel is considered in the model, while the pressure losses at the junctions of channels, which could be caused by abrupt changes in flow direction or channel dimension, are ignored. Thirdly, the fluid in the flow channel obeys ideal gas laws. Finally, the overlap area of channels at the junctions is neglected (which implies that the calculated total channel area is greater than the actual total channel area).For laminar flow, the pressure drop in a channel for a given flow rate can be determinedby:(cid:2)(cid:1)(1)whereis the channel length, is the hydraulic diameter, andis the fluid density,is the average flow velocity in the channel, is the friction factor.

2 Mesh of the flow field with × nodes (one inlet, one outlet), which were connected by -× × + ×channels.the There are varying volumetric flow rates through the channels. For a given node flow-in rate should equal the flow-out rate. The relationship can be expressed as a zero sum of all the flow rates through a node if the flow-in rate is defined as positive and the flow-out rate is negative, i.e.2.2.2 Constraints(2)(3)(4)(5)(6)(7)(8)The wider the channel is, the lower the pressure drop that occurs in the channel.However, the channel cannot be arbitrarily wide. The total surface area of the channel isThis area must be less than the area of the flow field. Surface area is an important parameter in the design of a bipolar plate flow field because it directly influences the mass transport efficiency and the electrical conductivity of bipolar plates.

Surface area is an important parameter in the design of a bipolar plate flow field because it directly influences the mass transport efficiency and the electrical conductivity of bipolar plates. Because the total flow field area equals the sum of the surface area of channels and the land area, which is the contacting area between bipolar plates and the gas diffusion layer (GDL), a larger surface area of channels would be good for mass transport but would reduce the conduction of electricity and eventually the fuel cell’s performance. Therefore, a trade-off between mass transport and electrical conductivity is needed in order to decide a proper channel surface area. In the present work, assuming that the surface area is constant, and taking 75% of the total flow field area as an example.

In the present work, assuming that the surface area is constant, and taking 75% of the total flow field area as an example. Then,whereis the flow field area of the fuel cell.2.2.3 Objective function, which is the current density when the reactant concentration decreases to zero at the surface of the is the highest current density that the fuel cell can operate at catalyst layer. In other words, under a certain condition. According to the electrochemical reaction and diffusion law, the limiting current density of a fuel cell [15] is:One important concept in PEM fuel cell design is the limiting current density,(cid:7)where ,is the number of moles of electrons transferred for one mole of reactant, e.g. for O2, is the is Faraday’s constant, is the thickness of the GDL.

for O2, is the is Faraday’s constant, is the thickness of the GDL. In order to improve the reactant concentration in the channel, and performance of a fuel cell, we need to increase the limiting current density [15] and thus the reactant concentration in the flow channel by optimizing the flow field design. Note that the reactant concentration decreases along the flow channel due to the reaction consumption and pressure drop from the inlet to the outlet. If the cathode is considered, the concentration of oxygen in the flow channel (assuming air is used) is:is the effective diffusivity within the GDL,(cid:2)(cid:3)is the pressure in where channels, is the absolute temperature. From equation (12), maintaining a high reactant concentration means maintaining a high and uniform pressure over the flow field, which in turn requires obtaining the lowest pressure drop over the flow field (if the same inlet pressure is used).

The known and unknown parameters are shown in Tables 1 and 2, can be treated as variables; respectively. In this model, both channel width however, for ease of design, only one of them is chosen as a variable, while the other is kept is chosen as a variable in our study, and a depth of 1.5 mm is constant [13]. Channel width used for all channels because channel width has a greater influence on the contact area of the reactant, the GDL and, consequently, the mass transport. The flow properties within channels with the same hydraulic diameter are almost the same; therefore, according to equation (6), channels with square cross-sections are used in the model first to simplify the calculation, where hydraulic diameter equals channel width from equation (2).

Then, the flow resistance for each channel can be determined from the hydraulic diameter equation (14), the pressure value at each node ( × unknowns) can be calculated by solving these × equations (16), and then the total pressure drop and the flow rate in each channel can be obtained from equations (13) and (15), respectively. The iteration process continues; at each iteration, the hydraulic diameter varies according to the scaling relationship given in equation (20), until the minimum pressure drop is achieved, thereby resulting in the optimum flow field design. A Matlab program was developed to implement this optimization algorithm.2.3 Optimized flow field designsSeveral optimal designs for different situations (i.e., positions of inlets and outlets) were obtained for a 5×5 cm2 flow field, with the total surface area of channels being 75% of the total flow field area.

The parameters, including the total flow rate, are given in Table 1 for known parameters and in Table 2 for unknown parameters. The results for several scenarios are shown in Fig. 3. For all scenarios, the inlet is located at the bottom left corner. Fig. 3(a) shows the initial mesh of the square flow field with 11×11 nodes and a pre-set channel width. Fig. 3(b) depicts an optimum result obtained from the proposed model; the pressure drop in the flow field is 24.3 Pa, and the outlet is located at the top right corner. In Fig. 3(c), the single outlet was changed to would uniform sinks over all the nodes. If the inlet flow rate is leave each channel to go into the GDL for reaction. In this optimum situation, all the reactant is consumed, and, moreover, the reactant is distributed uniformly to the GDL. Note that, realistically, the diffusion of reactant from the flow channel to the GDL occurs along the channels.

Note that, realistically, the diffusion of reactant from the flow channel to the GDL occurs along the channels. Here, we count only the total amount along one channel at the corresponding nodes. The optimized design of this situation is shown in Fig. 3(c), and the calculated pressure drop is 9.42 Pa. Along with the flow field designs including loops like those in Fig. 3(b) and (c), the designs without loops have also been investigated. Fig. 3(d) shows the design with 11×11 nodes on the flow field, and the pressure drop calculated is 6.02 Pa. Fig. 3(e) shows the design obtained from the flow field with 16×16 nodes, and the pressure drop calculated is 20.17 Pa. Both of these designs are based on a uniform-sink outlet. The above design examples show the capabilities of this basic mathematic algorithm, which obtains the desired flow field design with the desired, then the flow rate ofproperties from well-defined objective functions and constraints.

It can be used for flow fields of any shape and inlets and outlets at any locations. CAD models also were built based on the channel width data generated from the optimization model for FEM analysis and fabrication of bipolar plates with flow channels designed for experimental evaluation. Although the optimized designs have complex geometries with tapered channels, they can be fabricated easily with the Selective Laser Sintering (SLS) process [20, 21]. Fig. 4 shows the corresponding CAD models from Fig. 3(c) and (d).(a)(b) (c)(d) (e) Fig. 3 (a) Initial mesh of the flow field (unit: m); Optimum flow field designs obtained from the optimization model when (b) N = 11, one inlet with one sink; (c) N = 11, one inlet with uniform sink; (d) flow field without loops, N = 11, one inlet with uniform sinks; (e) flow field without loops, N = 16, one inlet with uniform sinks.

Finite element modeling (FEM) can be used to provide predictive models of PEM fuel cell operation [23, 24]. In our present work, an FEM model of a PEM fuel cell was developed with the commercial software FLUENT and used to investigate the performance of a fuel cell, including polarization curves, pressure distribution, and current density distribution within the fuel cell for a parallel design, a serpentine design, and an optimized flow field.The operation of PEMfuel cellsinvolves complicated processesTo verify that the FEM model established was correct and sufficiently accurate, the results from our FEM simulation were compared with the experimental results from reference [22], as shown in Fig. 5. The flow field was 5×4.92 cm2 with a serpentine design, the channel width and depth were 1.2 mm and 1 mm, respectively, and land width was 1.2 mm. The operation conditions were listed in Table 3. The parameters used in the FEM models were given in Table 4. Fig.

The operation conditions were listed in Table 3. The parameters used in the FEM models were given in Table 4. Fig. 5 shows that the FEM simulation results could closely predict the actual performance of PEM fuel cells, although the voltage and power density predicted by the FEM simulation were slightly higher than the experimental results due to neglecting some losses.Table 3 The operation conditions for comparing the FEM simulation with experimental results.Operation temperature (K) 323 Operation pressure (atm) 1 6.98 10-7 H2 flow rate (kg/s) 1.13 10-5 O2 flow rate (kg/s) Humidity 100% Table 4 Parameters used in the FEM simulation study.Reference current density at anode (A/cm2) Reference current density at cathode (A/cm2) GDL and catalyst layer porosity GDL and catalyst layer permeability (m-2) Exchange coefficient at anode Exchange coefficient at cathode Concentration exponent at anode Concentration exponent at cathode9.48×105 1.0×104 0.4 5.68×1010 0.5 1.5 0.5 1.0Fig.

5 Comparison of polarization curves and power density obtained from our FEM simulation and experimental results from [22].The FEM model was used to compare numerically the performance of the parallel (Fig. 4(c)), multiple-serpentine (Fig. 4(d)), and optimum designs (Fig. 4(a)) under the operation conditions given in Table 5. Fig. 6 shows the polarization curves and power density curves for the three different designs. Fig. 7 shows the pressure distribution in the flow channels on the cathode side and also the pressure distribution at the interface of the GDL and catalyst layer, where the operation voltage is 0.55 V. The corresponding oxygen distribution and current density distribution at the interface of the GDL and catalyst layer are shown in Figs.

The corresponding oxygen distribution and current density distribution at the interface of the GDL and catalyst layer are shown in Figs. 8 and 9, respectively.Table 5 Operation conditions used in the FEM simulation for different flow field designs.Operation temperature (K) Operation pressure (atm) H2 flow rate (kg/s) Air flow rate (kg/s) Humidity353 2.0 2.0 10-6 5.2 10-5 100%Fig. 6 Polarization curves and power density curves for the parallel design, the multiple- serpentine design, and the optimized design.Table 6 Pressure drop from inlet to outlet for different flow field designs (cathode side).Designs Pressure drop (Pa)Parallel design 226Multiple-serpentine design 1011Optimized design 2704. DiscussionFig. 6 clearly shows that the parallel design has the worst performance. This is expected because stagnant areas tend to form in the parallel channels, resulting in poor mass transport capabilities [4].

6 clearly shows that the parallel design has the worst performance. This is expected because stagnant areas tend to form in the parallel channels, resulting in poor mass transport capabilities [4]. This point is supported by the distribution of pressure and oxygen (Figs. 7 and 8). In some channels of the parallel flow field (e.g., channels 4 through 8 from left to right, Fig. 7(a, b)), the pressure difference is very small; therefore, no reactant is flowing through these channels and stagnant areas form. As shown in Fig. 8(a), no oxygen is supplied in these areas. The multiple-serpentine design performs similarly to the optimized design obtained from the proposed optimization model, when the fuel cells work below 2.2 A/cm2, and both of these designs produce almost the same maximum power density. However, considering the pressure drop over the flow field, the optimized design has a much lower pressure drop from the inlet to the outlet than the serpentine design. As seen in Fig.

However, considering the pressure drop over the flow field, the optimized design has a much lower pressure drop from the inlet to the outlet than the serpentine design. As seen in Fig. 7(c) and (e), and Table 6, the pressure drop is 270 Pa for the optimized design and 1011 Pa for the serpentine design. A lower pressure drop is better for the operation of PEM fuel cells and could reduce the need for an additional power supply, especially for a fuel cell with a large active area or a stack with a large number of fuel cell units. Fig. 7(a) shows that the pressure drop for the parallel design is only 226 Pa, which is even lower than the optimized design. This is because stagnant areas form in some channels of the parallel design, and the pressure drop in these channels is almost zero. For the optimized design, the pressure drop in every channel is required to be greater than zero in the optimization to satisfy the constraint that reactants flow through each channel at a certain flow rate.

Combining the results from polarization curves and pressure distribution, the optimized design isshown to have the advantages of both the serpentine design and the parallel design, performing as well as the serpentine design and having pressure drop as low as the parallel design. These merits of the optimized design result from the uniform distribution of pressure drops (Fig. 7 (e, f)) and the uniform transport of the reactant (Fig. 8 (c)).The polarization curves in Fig. 6 from the multiple-serpentine design and the optimized design are compared further, according to the different regions of polarization curve. Note that there are three regions for a typical fuel cell polarization curve corresponding to three different kinds of losses, i.e., active losses, ohmic losses, and mass transport losses [15]. The active losses (the first portion of the polarization curve) are almost the same for these two designs because their operation conditions are the same.

The active losses (the first portion of the polarization curve) are almost the same for these two designs because their operation conditions are the same. The two designs also show very similar ohmic losses. This can be seen from the middle portion of the polarization curve (i.e., the nearly linear portion). The ohmic losses are caused by the electrical (proton) resistances of the membrane, catalyst layer, GDL, and bipolar plates, and the contact resistances among them. All the resistances for the two designs are the same because the same materials are used, except the contact resistance between the GDL and the bipolar plates, which is mainly determined by the contact areas (land areas) between these two components and could be influenced by the design of flow field. In order to achieve fair comparison, in this study the flow fields are designed with the same land area, around 40% of the total flow field area. Therefore, the nearly same ohmic losses are obtained.

In order to achieve fair comparison, in this study the flow fields are designed with the same land area, around 40% of the total flow field area. Therefore, the nearly same ohmic losses are obtained. Note that although 75% of the total flow field area is used as the channel surface area in the optimization model, the actual channel surface area is only 59.2% because there are lots of overlapping areas for channels sharing the same nodes. In the optimization model, these overlapping areas are not eliminated for the purpose of algorithmic simplification. From the last portion of the polarization curves, the optimized design can operate at as high as 3.0 A/cm2 rather than 2.6 A/cm2 for the multiple-serpentine design, before the voltage drops to 0.3 V. Therefore, the optimized design suffers much less mass transport losses than the serpentine design, indicating that the optimized design has much more efficient mass transport capabilities.

9 Current density (unit: A/m2) distribution at the interface of the GDL and catalyst layer on the cathode side at 0.55 V for (a) parallel design, (b) multiple-serpentine design, and (c) optimized design.Comparing the pressure distribution of the three designs in Fig. 7, it is obvious that the optimized design has the most uniform pressure distribution over the whole flow field. Because the performance of a fuel cell is determined largely by the oxygen reduction reaction at the cathode, the distribution of oxygen is compared in Fig. 8. The catalyst surface of the optimized design has the largest area covered by oxygen, showing that it provides better mass transport of oxygen. The current density has a distribution (see Fig. 9) similar to oxygen because places receiving more oxygen generate more current. Note that the negative sign in the scale of Fig. 9 (b) and (c) indicates that the current is flowing in the reverse direction.

Note that the negative sign in the scale of Fig. 9 (b) and (c) indicates that the current is flowing in the reverse direction. Comparing these three designs in terms of performance, pressure distribution, oxygen distribution, and current density distribution leads to the conclusion that the optimized design provides power density as high as the serpentine design but with much lower pressure drop, more uniform pressure distribution, and more efficient mass transport capability. The optimized design also allows the fuel cell to work at a higher current density. Therefore, the proposed optimization model is suitable for the design of flow fields for bipolar plates and the improvement of PEM fuel cell performance.5. ConclusionsAn optimization model has been proposed for the design of flow fields in order to achieve a lower pressure drop in bipolar plates and higher PEM fuel cell performance.

ConclusionsAn optimization model has been proposed for the design of flow fields in order to achieve a lower pressure drop in bipolar plates and higher PEM fuel cell performance. The optimization model generates flow field designs similar to structures found in nature (i.e., leaf veins). It also demonstrates this model’s ability to optimize designs for any shapes of flow fields, any locations of inlets and outlets, and any desired flow properties. The comparison of FEM simulation results between the optimized design and the conventional parallel and multiple- serpentine designs has demonstrated that the optimized design can provide power density as high as the serpentine design but with much less pressure drop, more uniform pressure distribution, and more efficient mass transport capability.

The optimized design also allows the fuel cell to work at a higher current density.AcknowledgementThis project is supported by Air Force Research Laboratory under contract #FA8650-04-C-5704.References[1] Blunk, R., Elhamid, M. H., Lisi, D., Mikhail, Y. Polymeric composite bipolar plates forvehicle application. Journal of Power Sources, 2006, 156: 151-157. [2] Watkins, D. S., Dircks, K. W., Epp, D. G. US Patent No. 5,108,849, 1992. [3] Watkins, D. S., Dircks, K. W., Epp, D. G. US Patent No. 4,988,583, 1991. [4] Li, X., Sabir, I. Review of bipolar plates in PEM fuel cells: Flow-field designs. InternationalJournal of Hydrogen Energy, 2005, 30: 359-371.[5] Li, X., Sabir, I., Park, J. A flow channel design procedure for PEM fuel cells with effectivewater removal. Journal of Power Sources, 2007, 163: 933-942.[6] Misran, E., Daud, W. R. W., Majlan, E. H. Review on serpentine flow field design for PEMfuel cell system.

ocompatibility of SLS-Formed Calcium Phosphate ImplantsG. Lee*, lW. Barlow*, w.c. Fox**, and T.B. Aufdennorte***Department of Chemical Engineering The University of Texas at Austin Austin, Texas 78712**BioMedical Enterprises, Inc. 14785 Omicron Dr., Suite 205 San Antonio, Texas 78345-3201***Department of Pathology The University of Texas Health Science Center at San Antonio San Antonio, Texas 78284AbstractA method for fabricating artificial calcium phosphate bone implants by the Selective Laser Sintering (SLS) process has been developed that can fabricate complex and delicate calcium phosphate bone facsimiles from a variety of data inputs including Computed Tomography(CT) files (1). This paper discusses two in vivo biocompatibility studies of SLS-formed calcium phosphate implants in both rabbits and dogs.

This paper discusses two in vivo biocompatibility studies of SLS-formed calcium phosphate implants in both rabbits and dogs. Histologic analysis shows a high degree of biocompatibility and bone ingrowth in both studies.IntroductionMuch attention has been given to develop materials that can assist in the regeneration of bone defects and injuries (2). in developing a biocompatible material, little research has been done toward the rapid and automated shaping of replacement bones using these materials. A hwnan body is composed of 206 different bones and every bone is different in size from person to person. Specially, bones in craniofacial or maxillofacial area are prone to be very complex in shape as well as vary in size.

Specially, bones in craniofacial or maxillofacial area are prone to be very complex in shape as well as vary in size. For example, reconstruction of the oral cavity could benefit from an implant that matched the shape of the patient's original anatomy, fit the defect and guided the regeneration of the soft tissue.Although efforts has been greatAs described earlier, we developed a process that uses SLS to fonn complex geometries from calcium phosphate powders (1). This process, coupled with the well established medical imaging technology of Magnetic Resonance Imaging (MRI) and CT, can precisely reproduce accurate facsimiles of any skeletal bone. The present study examines the biocompatibility of implants made by this process in rabbit and dog animal models.15Materials and MethodsPowder preparation Monocalcium phosphate, monohydrate (MCP, Food grade) and Dicalcium Phosphate Anhydrous(DCP/A, USPIFCC grade) were obtained from Rhone Poulenc Basic Chemical Co. as raw materials.

as raw materials. Specimens with CaIP ratio of 0.5 to 0.9 were forrned in a two step process. The first step involved the mechanical mixing of MCP and DCP/A in such a way that the resulting powder contained a CalP ratio of 0.9. For example, 0.8 mole of DCP/A and 0.2 mole ofMCP were mechanically mixed and fired at a controlled temperature about 950°C to obtain a solid solution of l3-calcium pyrophosphate (3), 13-2CaO.P205. The resultant sintered cake was then either ball milled or ground by mechanical blender, and selectively sieved to obtain powders with particle size less than 75 I-lmFollowing previously established procedures (4), a poly(methymethacrylate-co-n-butyl methacrylate) latex polymer was synthesized as a binder. The melt flow index of polymer was tailored about 7.7 gj10 min. at 200°C and 75 psi, as measured by a Kayness Galaxy I capillary rheometer.

The melt flow index of polymer was tailored about 7.7 gj10 min. at 200°C and 75 psi, as measured by a Kayness Galaxy I capillary rheometer. This flow range had been previously discovered to allow the binder to flow and rapidly wet the inorganic material under SLS process conditions.Calcium phosphate powder was coated with polymer by spray drying a slurry of particulate and emulsion binder. The spray drier used in this study is a pilot -plant scale Anhydro Laboratory Spray Drier # 1 with co-current rotary atomizer. This spray drying process produces a free flowing powder that has the desired agglomerated morphology. Spray drying condition were as follows. Solids content of slurry was 45 wt. %, inlet temperature, 175°C, outlet temperature, 110°C, and atomizer speed was above 30,000 Polymer content in coated powder was 22 wt. % as deterrnined by Thermal rpm.

%, inlet temperature, 175°C, outlet temperature, 110°C, and atomizer speed was above 30,000 Polymer content in coated powder was 22 wt. % as deterrnined by Thermal rpm. Gravimetric Analysis (TGA).SLS processing Spray dried material were processed using a beta test workstation SLSTM Model 125 equipped with modulated 25 W CO2 laser. Laser output power, beam scan speed, scan spacing and powder bed temperature were controlled to optimize the green strength according to the shape of the implants. Optimal settings for Energy Density (5) were found to be in the range of 1 to 1.5 cal/cm2.Post Processing SLS-formed green objects were infiltrated with calcium phosphate solution, fired to remove the polymeric binder, then fired at higher temperature to sinter The parts were subsequently infiltrated with DCPIA the calcium phosphate particles. solutions in diluted phosphoric acid by submersing them in the solutions.

solutions in diluted phosphoric acid by submersing them in the solutions. The capillary action of the pores drew the infiltration solution into the part. After the pores were filled with infiltrant, the parts were allowed to dry prior to firing. The firing temperatures were controlled to be lower than the initial sintering temperature to avoid further shrinkage, but high enough to provide atomic diffusion and chemical reactions to occur between Ca and P in the infiltrant and that of the green part. When the green part was first infiltrated, the CaIP ratio and solids content of the infiltrant were chosen to facilitate a clean bum out of polymer. After debinding, the extent of infiltration and the concentration of the infiltrant16were varied to achieve desired density, pore structure, mechanical strength and a final CaIP ratio.

After debinding, the extent of infiltration and the concentration of the infiltrant16were varied to achieve desired density, pore structure, mechanical strength and a final CaIP ratio. This approach allowed the density of the final product and its chemistly to be precisely controlled in the final firing step.Material characterization The effect of post processing on calcium phosphate chemistly The microstructure of the prepared specimen was was studied by X-ray diffraction. evaluated by mercury porosimetry and scanning electron microscopy (SEM). The SEM was primarily used to visualize surface morphology and the internal structure of samples.The pore structure of the in vivo oral implant, described below, was measured by mercury porosimetry (Poresizer 9320, Micromeritics). Prior to porosimetry measurements, the oral implant was heated to expel the absorbed moisture and then stored in a desiccator until analyzed. The sample in the Poresizer was initially evacuated to 25 /-lm Hg.

The sample in the Poresizer was initially evacuated to 25 /-lm Hg. Mercury was then admitted and pressure was increased up to 30,000 psi. Assuming a contact angle of 130 0 (6), pore diameters ranging from 200 to 0.006 !Jm were measured. Figure 1 shows a typical pore size distribution. From mercury intrusion analysis, the total pore area 1.82 m2/g and median pore diameter 33 !Jm were measured. The measured porosity was about 30 %, which is smaller than the observed value, about 50 %, from geometrically based density measurements This could indicate that some closed pores are present in the part.o.V-o .0o .0o .0o .0o .0o. .~,ft.,I I'I \ i ! I ! i : I i ! i I I !1\ ! \Io.~I./ I..... 100\\ .....A-_.0.01Diameter (micron)Figure 1. Pore size distribution of oral implantTwo test methods, compression and 4-point bending, were used to quantify the strength of the specimens by a MTS Sintech liS material test frame.

Pore size distribution of oral implantTwo test methods, compression and 4-point bending, were used to quantify the strength of the specimens by a MTS Sintech liS material test frame. Typical strengths are in the range 2,000 psi in compression and 2,200 psi in 4-point bending.Implant preparation and animal studies. Preliminary in vivo biocompatibility studies were carried out with rabbits. Tiny rectangles, 0.12" x 0.16" x 0.05", were prepared by a process that simulated SLS process conditions. Calcium metaphosphate was mixed with17the same polymeric binder used previously (1) and poured into the cavity of the mold without compaction. The mold was heated to form polymer-bound "green" parts. The green parts were then fired at about 920°C to remove the binder and to sinter the calcium phosphate. One of the prepared calcium phosphate specimens was implanted over the frontal lobe of the skull in each of four New Zealand White Rabbits.

One of the prepared calcium phosphate specimens was implanted over the frontal lobe of the skull in each of four New Zealand White Rabbits. The cortex of each skull was decorticated to cause minor hemorrhage and create a small defect in which the implant could rest. At twelve weeks the animals were sacrificed. Implants and bone were retrieved and embedded in plastic, sectioned and analyzed.More elaborate in vivo biocompatibility studies were carried out in dogs. Anatomical measurements of the original geometry of the dog alveolar ridges were taken to determine their anatomy. AutoCad engineering computer-aided design software was then used to design the oral implants. These implants were rectangular shaped and tapered so as to decrease their buccal-lingual width apically, see Figure 2. The sides and bottom of these implants had large hexagonal macropores of about 2.4mm that penetrated entirely through the device. These macropores were aligned so they intersected within the implants.

These macropores were aligned so they intersected within the implants. Tooth extraction of the first mandibular molar bilaterally was performed in six dogs and bone was cut from the extraction site to prepare a defect in the alveolar ridge. The first molar was extracted bilaterally from the mandible. Primary closure was achieved and the dogs were allowed to heal for eight days. After the 8-day healing period, the extraction defect in the alveolar ridge was surgically exposed. Implants were placed in the ridge and primary closure was achieved. Dogs were put on soft food for the duration of the study. Radiography was performed to obtain data giving an indication of the implants osteoconductive properties and its biocompatibility. Radiographs were taken immediately post-operatively and at four weeks into the study.Figure 2.

Radiographs were taken immediately post-operatively and at four weeks into the study.Figure 2. Designed oral implantResults and DiscussionA histologic section of the calcium phosphate implant after twelve weeks implantation in rabbit's frontal skull is shown in Figure 3. The implant shows a high degree of biocompatibility and bone ingrowth. No evidence of foreign body giant cell or macrophage response was found in or adjacent to the specimen. There was no acute or chronic inflammation, and a high degree of bone apposition and infiltration is observed. Mineralized new bone is present throughout a significant portion of the implant closest to the underlying calvaria. Osteoid, or a collagen matrix with bone forming potential, fills It is suggested from this histology that, in time, the majority of the rest of the implant.18the specimen could be completely filled with dense mineralized bone. A higher magnification of the histologic section is shown in Figure 4.

A higher magnification of the histologic section is shown in Figure 4. Direct bone apposition and Host bone is ingrowth are visible across the interface and throughout the implant. directly connected to bone within the implant providing good mechanical stability.Figure 3. Histologic section of implants taken overlying the rabbit calvaria. (40x original magnification)Figure 4. Histologic section of implant taken overlying the rabbit calvaria (100x original magnification)A scanning electron micrograph of surface of the SLS fonned oral implant is shown in Figure 5. This surface is imaged at 150 times of the original magnification and 100 Jlm In Figure 6, the polished surface is measure is visible in the lower right of the image. It is clear from these pictures and mercury intrusion shown at the same magnification. analysis, that some of the pores are larger than 100 Jlm and that the pores are open and well interconnected.

It is clear from these pictures and mercury intrusion shown at the same magnification. analysis, that some of the pores are larger than 100 Jlm and that the pores are open and well interconnected. These issues are critically important to the successful development of a osteoconductive bone implant material. It is known that mineralized bone growth into porous implants required a minimum interconnected pore size of 100JlID, and the pore size showing the potential for the ingrowth of osteoid tissue, was found to lie between 40 and 100Jlm (7).Figure 5. Scanning electron micrograph of the surface of SLS-fonned implant.Figure 6. Scanning electron micrograph of a polished surface of SLS-fonned implant.19Figure 7 shows the periapical radiograph of a SLS-formed alveolar implant. This image was taken immediately post-operative. The rectangular form of the implant and hexagonal form of the macropores are visible in this radiograph.

This image was taken immediately post-operative. The rectangular form of the implant and hexagonal form of the macropores are visible in this radiograph. In Figure 8, a radiograph image taken four weeks following placement of the implant is shown. Crestal resorption and bone infiltration into the apical macropores and apical portion of the crestal macropore are evident from the film.Figure 7. A radiograph of a SLS fonned alveolar implant with micro and macropores.Figure 8. A radiograph of a SLS fonned alveolar implant with micro and macropores.As a demonstration of this bone fabrication technique, a human anatomical facsimile was formed from CT images using SLS.

A radiograph of a SLS fonned alveolar implant with micro and macropores.As a demonstration of this bone fabrication technique, a human anatomical facsimile was formed from CT images using SLS. This inferior view of the maxilla shows some dentition and anterior nerve and vascular foremen with great'resolution.Figure 9: CaIP human maxilla formed -from CAT images using SLS.20ConclusionBased on the excellent biocompatibility of material and the great extent of bony tissue ingrowth shown in the rabbit in vivo study, the calcium phosphate powder system for the SLS process has been established. As a starting material CaJP ratio 0.9 was synthesized and formed by SLS. Oral implants in alveolar ridges of dogs show excellent integration with host bone after only four weeks of implantation.

As a starting material CaJP ratio 0.9 was synthesized and formed by SLS. Oral implants in alveolar ridges of dogs show excellent integration with host bone after only four weeks of implantation. This study shows that a porous, resorbable calcium phosphate sYnthetic bone material can be precisely fabricated with complex geometry and successfully used in vivo for filling and replacing defective bone.AcknowledgmentsThe authors gratefully acknowledge partial support for this work by the State of Texas Advanced Technology Program, No. 265, BioMedical Enterprises, Inc., and the United States Air Force Armstrong Laboratory through contract F41624-95-C-2008.References1G. Lee and J.W. Barlow, "Selective Laser Sintering of Calcium Phosphate Powders," Solid Freejorm Fabrication Symposium Proceedings, 1993, The University of Texas at Austin, Texas, pp.191-197.2.S.F. Hulbert, L.L. Hench, D.Forbes, and L.S. Bowman, "History of Bioceramics," Ceramics in Surgery edited by P.

Hulbert, L.L. Hench, D.Forbes, and L.S. Bowman, "History of Bioceramics," Ceramics in Surgery edited by P. Vincenzini, 1982, Elservier, Amsterdam, pp 3-29.3. W.L.Hill, G.T.Faust and D.S. Reynolds, "The Binary System P205-2CaO.P205.", Am. Jour. Sci., 1944,242 (9),457-477.4. N.K.Vail, et aI., "Develpment of a Poly(methyl methacrylate-co-n-butyl methacrylate) Copolmer Binder System", J. Appl. Polym. Sci., 1994,52, 789-812.5.I C. Nelson, et ai, "Selective Laser Sintering of Polymer-Coated Silicon Carbide Powders," Ind. Eng. Chern. Res., 1995,34,1641-1651.6. E.W. Washburn, "Note on a Method of Deterrnining the Distribution of Pore Sizes in a Porous Material," Proc. Natl. Acad. Sci., 1921, 7, ppI15-116.7. I I Klawitter and S.F. Hulbert, " Application of Porous Ceramics for the Attachment of Load Bearing Internal Orthopedic Applications," I Biomed. Mater. Res. Symposium, No.2, 1970, ppI61-229.2122.

Biodegradable Polycaprolactone based ink developed for 3D Ink Jetting Yinfeng He1, Sam Kilsby2, Chris Tuck1, Ricky Wildman1, Steven Christie2, Hongyi Yang1, Steven Edmondson3, 1Additive Manufacturing and 3D Printing Research Group, Nottingham University, UK 2Department of Chemistry, Loughborough University, UK 3Materials Department, Manchester University, UK epxyh@exmail.nottingham.ac.ukAbstract: Biomedical applications are one of the driving forces for Additive Manufacturing, however to extend the range of applications and markets new materials are required. A new type of biodegradable Polycaprolactone (PCL) based ink that is suitable for 3D inkjet printing was successfully developed. UV curable PCL was synthesized and mixed with Poly(ethylene glycol) di-acrylated (PEGDA) to prepare an ink with suitable viscosity for inkjet printing. Their mechanical properties as well as the printing accuracy were measured by nano-indentation and scanning electron microscopy.

Their mechanical properties as well as the printing accuracy were measured by nano-indentation and scanning electron microscopy. Post curing was applied to printed samples in order to study how post curing may influence sample properties. It was found that within 30min post-curing period, the sample’s surface which is direct illuminated by UV light increased from 31.22MPa to 70.20MPa while the bottom surface showed less incensement from 34.9MPa to 39.8MPa.Introduction3D Printing or Additive Manufacturing (AM), as a disruptive manufacturing technology, has been attracting increasing attention in recent years. Many studies have been carried out to enable AM to become a process that is able to manufacture end-use products. One of the potential applications for AM is in the making of biodegradable or bioresorbable medical products with the potential of tailored drug release functions.

One of the potential applications for AM is in the making of biodegradable or bioresorbable medical products with the potential of tailored drug release functions. Some biomedical products require complex and strict pore structures to achieve enhanced cell adhesion and growth [1-3]. These requirements in pore structure and sizes are difficult to be controlled with current manufacturing techniques (such as foaming). However with AM technology, the sizes and locations of each pore can be accurately controlled and this lends the possibility of being able to manufacture bespoke products.Polycaprolactone (PCL) is widely used in biomedicine as it can gradually decompose by random hydrolytic chain scission of the ester groups [4-5]. Some researchers have attempted to build PCL structures with additive manufacturing methods such as Powder Bed Fusion (i.e. laser sintering) or material extrusion [6-8]. However, little work has been done in processing PCL by 3D inkjet printing.

laser sintering) or material extrusion [6-8]. However, little work has been done in processing PCL by 3D inkjet printing. Compared with other AM techniques, 3D inkjet printing has the potential of producing multi-material artefacts within a single process cycle with controlled material distribution. This facility can be used to manufacture products with localized drug distribution or degradation speeds. Therefore, developing a biodegradable material which is1052suitable for 3D inkjet printing could bring the manufacturing of biomedical product into a brand new field.3D inkjet printing requires a low viscosity ink, which solidifies quickly after deposition hence photo-curing has been a popular method of achieving fast solidification for commercial materials. Although pure caprolactone does not cure under UV irradiation, groups (normally acrylated group) can be grafted onto the end of a PCL polymer chain to make it UV curable [9-10].

Although pure caprolactone does not cure under UV irradiation, groups (normally acrylated group) can be grafted onto the end of a PCL polymer chain to make it UV curable [9-10]. As printers have very strict requirements on ink viscosity, a diluent is also a necessary component to help adjust the inks viscosity. PolyEthyleneGlycol (diacrylated) (PEGDA), can be used as such a diluent, and has been widely used for UV curable biocompatible materials [11-13]. Some authors have suggested that a copolymer of PCL and PEG can help develop products with different surface function and modify the drug release profile of a material because PCL is hydrophilic and PEG is hydrophobic [14-15].In this paper, the printability of UV curable PCL: Polycaprolactone di-methacrylated (PCLDMA) was assessed. Rheological data were collected with different PEGDA proportions in a temperature range from ambient temperature to 60ºC to help select suitable processing conditions and PEGDA concentration for printing.

The synthesised UV curable PCL’s viscosity was adjusted by adding 30wt% of PEGDA and successfully printed by using a Dimatix DMP2800 inkjet printer.3D structures were then created with this ink and characterized to help understand the properties of this material.MethodologyInk preparationPCLDMA (synthesised) and PEGDA (Sigma-Aldrich average Mn~250) were added into an 8ml amber vial and stirred at room temperature for 15mins at 800rpm using an IKA RCT Basic IKAMAG Magnetic Stirrer (with Temperature Controller). 3wt% of photo-initiator (2,4-Diethyl-9H-thioxanthen-9-one(DETX),sigma-aldrich,98%) and 3wt% of accelerator (Ethyl 4-(dimethylamino)benzoate(EDB),sigma-aldrich,99wt%) were added into the PCLDMA:PEGDA mix and stirred at 85ºC for 5mins until all the solutes are fully dissolved. Before printing, the prepared ink required a degassing procedure to remove dissolved oxygen and to help minimize the 'oxygen inhibition' effect [17-18].

Before printing, the prepared ink required a degassing procedure to remove dissolved oxygen and to help minimize the 'oxygen inhibition' effect [17-18]. The degassing procedure was carried out by purging the mixed ink with nitrogen gas for 15minutes. This procedure created lots of Nitrogen bubbles within the ink which seriously reduces the droplet formation stability of the ink. Therefore the ink was prepared 24hrs prior to printing; the degassed ink was then settled to release the bubbles.Printability Assessment1053The key parameters for determining ‘’printability’’ are viscosity and surface tension. Normally, viscosity is the most fundamental parameter used to decide whether an ink is printable or not. The printing viscosity range varies dependent on the printheads being used. Surface tension should also be taken into consideration. It has been reported that the inverse Ohnesorge number (Oh-1) can be used as the printing indicator (Z) to help predict an ink’s printability.

Surface tension should also be taken into consideration. It has been reported that the inverse Ohnesorge number (Oh-1) can be used as the printing indicator (Z) to help predict an ink’s printability. Z takes both viscosity and surface tension into consideration and is shown following equation[16].√Where is density, is characteristic length, is surface tension of the fluid and is viscosity of ink.The viscosity of the PCLDMA:PEGDA mixture (under shear rates of 100s-1 and 1000s-1) was measured by a cone plate rheometer (Malvern Kinexus Pro) to identify the PCLDMA:PEGDA proportion and the processing temperature that would give a suitable viscosity for inkjet printing. Each measurement started at 25ºC with 5º C increments up to 60ºC. A protocol of waiting 300s after reaching the test temperature was set to ensure the ink was in a steady state condition. At each temperature point and shear rate, the viscosity was recorded at 5s intervals within a 180s test time..

At each temperature point and shear rate, the viscosity was recorded at 5s intervals within a 180s test time.. Surface tension was measured by a pendant drop method using a Kruss DSA100S. The shape was captured at the equilibrium state and used to calculate the surface tension.Sample properties assessmentThe ink with the final composition was then injected into a print cartridge. The injection procedure was carried out in the dark to prevent light irradiation and careful attention was paid to avoid bubble formation within the ink. The cartridge was wrapped with foil tape to make sure the ink was not cured inside the cartridge by ambient light. About 2ml of the prepared ink was injected into a disposable cartridge and printed by DimatixDMP-2800 material printer. The printed ink was cured by real- time UV curing; a UV curing unit was mounted directly on to the printing unit (Figure 1) to move in conjunction with the print direction.

The printed ink was cured by real- time UV curing; a UV curing unit was mounted directly on to the printing unit (Figure 1) to move in conjunction with the print direction. A further UV LED unit (intensity of 1000mW/cm2) was used to examine the influence of different post-curing time (10 minutes, 20 minutes and 30 minutes) on the mechanical properties of the printed parts.The mechanical properties of printed sample were characterized by nano-indentation at room temperature (Micro Materials, NanoTest NTX with hot stage and inert gas cabinet). Both the top and bottom surface were characterized. Load-depth curves were recorded on 5×5 grid with 100µm separation between each indentation. The1054applying force was set to 5mN with a 0.25mN/s loading and unloading rate and a spherical indenter with 50µm radius was used. The Hardness is calculated by:Where is applied load, is the radius of the circle of contact.

The Hardness is calculated by:Where is applied load, is the radius of the circle of contact. The radius of the circle of contact is calculated by follow equation:√ ( ) ( )Where R is the radius of the spherical indenter,, is the total penetration depth, is the residual depth and c is correction constant for piling-up or sinking-in effect [19],The indentation modulus E can be calculated by:Where P and are applied load and radius of circle of contact respectively, is the elastic deformation depth.Mesh structures were printed to help understand the manufacturing accuracy that PCLDMA:PEGDA (70:30) ink could achieve. Printed mesh structures were sputter coated with Platinum and examined by SEM (XL30 ESEM Philips).Results and DiscussionPrintability AssessmentThe viscosity results of PCLDMA with different PEGDA proportions were measured (Table 1).

As PEGDA is a diluent and only biocompatible after curing, when choosing the composition, the amount of PEGDA needs to be as low as possible within the printable viscosity range, in order to maximise the biodegradability of final product. PCLDMA with 30wt% of PEGDA at 60ºC was chosen as the final proportion in this paper based on the rheological tests and printing requirements. However, as different printheads have different printing viscosity ranges, this does not apply to all the piezo based printheads. The viscosity distributions of PCLDMA:PEGDA mixture with various proportions under different environment temperatures are given in Table 1, which can help decide the optimum composition for the other printheads.As PEGDA is only bio-compatible but not biodegradable material, it will reduce the biodegradability of the final product. Therefore, when choosing the final composition, the one with higher PCLDMA concentration but still within printable range is1055preferred.

Therefore, when choosing the final composition, the one with higher PCLDMA concentration but still within printable range is1055preferred. Based on this principle, PCLDMA:PEGDA (70:30) was chosen as the final composition, which will be used for following printing.The printability indicator, Z, for PCLDMA:PEGDA (70:30) was then calculated (Table 3). It has been suggested that when the value of the printing indicator is between 1 and 10, the ink will normally be printable [16]. From Table 2, it can be seen that based on these calculations, the Z values of PCLDMA:PEGDA (70:30) at 60 ºC were in the printable range.The viscosity of the ink was measured throughout the whole ink preparation procedure to monitor viscosity variations (Table 3). It can be seen that after adding photo-initiators and being degassed, the viscosity of the ink increased by 5-10%. This was due to two reasons: the adding of solid content and a small amount of curing during the degassing procedure.

This was due to two reasons: the adding of solid content and a small amount of curing during the degassing procedure. As one may expect, adding high viscosity content will increase the viscosity of the ink. The photo-initiator and accelerator used in our experiment are both solids and they occupied 6wt% of the whole ink which led to viscosity increase. Also, the ink became quite reactive during degassing and as the degassing procedure was not carried out in a completely dark environment there will be small amount of curing which also increased the prepared ink’s viscosity.Real-time curing and Post-curing effectFive square specimens, 5mm (W)*5mm (L), were prepared for nano-indentation testing. The printing pattern and samples are shown in Figure 3 where 100 layers were printed. The final thicknesses of these square samples were ~500μm.The hardness and indentation modulus of printed samples with different post curing time were measured by nano-indentation (Table 4).

The final thicknesses of these square samples were ~500μm.The hardness and indentation modulus of printed samples with different post curing time were measured by nano-indentation (Table 4). The measurements were carried out on the sample’s top and bottom surface respectively. As the samples were produced by stacking up layers of material, those layers printed in the early stages will inevitably be repeatedly exposed to UV illumination when following layers are printed. This will lead to a printed sample which has a gradient of UV exposure time from bottom to top. As UV exposure time is normally related to curing level, this may eventually result in property deviation and therefore, both surfaces were measured separately.The indentation results (Table 4) showed the top and bottom surface of the sample has very similar properties before any post-curing treatment.

The hardness and indentation modulus on both surface have variations however these are within the testing deviation range.The mechanical properties of the sample’s top surface, which was directly illuminated by UV light, had a significant increase after 10mins of post-curing (Table 4 and Figure 4).However additional post-curing for 20mins and 30mins did not further influence these properties. The rise of hardness and modulus on the top surface after post-curing was mainly due to and increase of cross-link density. Prior work has shown that the hardness and modulus of a crosslink material is positive in relation to its crosslink density [20, 21]. At low crosslink density, polymer chains are1056less restricted, therefore it can easily deform with an applied manifesting as a low hardness and modulus. As crosslink density increases those free segments are connected with each other building an increasingly dense network.

As crosslink density increases those free segments are connected with each other building an increasingly dense network. The mobility of the polymer chain segments becomes restricted and the specimen will then show stronger resistance to an applied force. When a specimen was printed, the conversion of the C=C group into the covalent crosslink cannot normally reach 100%. During the post-curing procedure UV illumination will provide extra energy to help the residual C=C group form new crosslink and therefore, further increase the crosslink density and hence its hardness and modulus.Meanwhile, the properties of the sample’s bottom surface did not show a notable change with the increase of post-curing time. This could be because during the post- curing procedure, samples were illuminated from the top surface and the UV irradiation needs to penetrate the whole sample before reaching the bottom surface.

This could be because during the post- curing procedure, samples were illuminated from the top surface and the UV irradiation needs to penetrate the whole sample before reaching the bottom surface. During the penetration procedure, the intensity of the UV light would be reduced by absorption from the sample and result in only a small quantity of radiation reaching the bottom surface. From the nano-indentation results, the bottom surface did not receive enough energy to achieve further crosslinking and therefore the properties remain unchanged. UV absorbance spectrums with a different thickness of cured ink films are currently under characterization to affirm this hypothesis and also help discover the penetration depth.Printing and CharacterizationMesh structures were then printed and a schematic figure of the printing pattern was shown in Figure 5 (a). Meshes with three different wall thicknesses (150μm, 300μm and 500μm) were printed onto glass slides for futher SEM examination.

Meshes with three different wall thicknesses (150μm, 300μm and 500μm) were printed onto glass slides for futher SEM examination. The distance between each wall was set as 1mm to allow each printed vertical or horizontal wall to be seperated from each other. Ten layers of PCLDMA:PEGDA (70:30) were printed and the sample appearance is shown in Figure 6(b).The mesh structures were then observed under SEM (Figure 6). The actual printed wall thickness was measured and calculated to help analysing the dimensional variation between the printed structure and the original design. From the results in Table 5, it can be noticed that the printed mesh structure with larger designed wall thickness had less deviation percentage compared with the meshes with thinner walls.

From the results in Table 5, it can be noticed that the printed mesh structure with larger designed wall thickness had less deviation percentage compared with the meshes with thinner walls. However the deviation was always around 40μm which would indicate that the processing accuracy of Dimatix material printer for PCLDMA:PEGDA (70:30) is around 40μm, when using printhead with 21μm nozzles.Figure 6 also showed that under the print conditions the PCLDMA:PEGDA (70:30) ink could not form accurate and sharp edges. Rectangular gaps were designed inside the mesh structure. However in the actual printed structure, the gap morphed into rounded rectangular shapes. Meanwhile, dislocation of printed ink droplets can be observed from the SEM pictures and causing rounded rectangular gaps as well1057as curving walls. These might be due to the slow curing speed of the printed ink.

These might be due to the slow curing speed of the printed ink. Although, the ink was illuminated by UV light immediately after being printed, the illumination time was quite limited in a single scan. This is becausethe UV curing unit was attached and moving with the printhead. So the energy provided in a single scanning may not be enough to allow freshly deposited ink to become fully cured immediately. Therefore, merging and dislocation of the uncured ink droplet could happen due to gravity, movement of the platform and merging with the ink deposited in subsequent printing cycles. All these will lead to rounded edge and curing walls.However, as the illumination area of UV curing unit were larger than the printed area of each printing cycle, the previously printed ink can still receive UV illumination during the following printing. So the ink will receive discontinuous UV illumination and finally be cured after obtaining enough energy.

So the ink will receive discontinuous UV illumination and finally be cured after obtaining enough energy. But the curing time will be enlarged compared with continuous UV illumination. Increasing the curing speed could help printed ink cure in a shorter period of time, hence reduce the chance of merging and dislocation happening, improving the print quality. This could be achieved by either increasing the intensity of UV illumination or creatingan oxygen free environment.Figure 6 (d) is a SEM image showing a cured mesh surface in high magnification. Wrinkles (about 1 to 2μm) can be observed on the entire printed mesh structures’s surface. A similar self-wrinkling effect was observed by Chandra et al [22]. They suggested that this effect was mainly due to oxygen inhibition which caused crosslinking speed variation from the top to the bottom.

They suggested that this effect was mainly due to oxygen inhibition which caused crosslinking speed variation from the top to the bottom. When UV curable films were exposed to UV illumination within the presence of oxygen, a thin layer at the top surface will remain uncured due to the oxygen inhibition. A crosslink gradient would be formed through the depth direction because of oxygen concentration gradient formed at the surface by diffusion. This situation will lead to in-plane stress and cause surface wrinkle. In Chandra et al.’s work, they also concluded that by controlling the oxygen concentration in the environment, the size of surface wrinkle could also be controlled.Figure 7 is a printed curving mesh structure with PCLDMA: PEGDA (70:30) ink. 50 layers were printed and surface profiling data Figure 7(d) showed the total height of the structure was around 250μm. Figure 8 shows optical microscopy images of a printed curving mesh structure.

50 layers were printed and surface profiling data Figure 7(d) showed the total height of the structure was around 250μm. Figure 8 shows optical microscopy images of a printed curving mesh structure. Similar effects were also observed that the ink did not fully cure immediately after deposition and droplets at the edges falling down to the substrate forming coarse structures at the base.ConclusionA PCLDMA: PEGDA ink that is suitable for 3D inkjet printing to produce biodegradable 3D structures has been demonstrated for the first time. For different printers, the proportion of PEGDA and processing temperature can be varied, which should be decided based on the given rheology database.1058In this paper, PCLDMA: PEGDA (70:30) was chosen and observed to be suitable for a Dimatix DMP-2800 when printed at 60ºC . The prepared ink can be cured sufficiently to retain expected structures during printing and stable products can be produced.

The prepared ink can be cured sufficiently to retain expected structures during printing and stable products can be produced. The hardness of printed samples was around 5MPa with an indentation modulus of 30MPa. These properties increased when a post-curing procedure was applied. However, only the mechanical properties at the top surface were improved. From SEM examination, it was found that print quality was influenced by curing speed and wrinkles were observed on the surface of the cured structures. For future work, the curing efficiency and printing qualities of PCLDMA:PEGDA (70:30) ink with different photo-initiator / accelerator ratios should be investigated.

bmitted to: SFF Austin 2004 By Dr Wei SunBiomimetic Design and Fabrication of Interior Architecture of Tissue scaffolds using Solid Freeform FabricationB. Starly, A. Lau, W. Sun Dept of Mechanical Engineering and Mechanics Drexel University Philadelphia, PA 19104andW. Lau, T. Bradbury Therics, Inc. 115 Campus Drive Princeton, NJ 08540Abstract: Modeling, design and fabrication of tissue scaffolds with intricate architecture, porosity and pore size for desired tissue properties presents a challenge in tissue engineering. This paper will present the details of our development in designing and fabrication of the interior architecture of scaffolds using a novel design approach. The Interior Architecture Design (IAD) approach seeks to generate scaffold layered freeform fabrication tool path without forming complicated 3D CAD scaffold models.

IntroductionThe loss or failure of an organ or tissue is one of the most frequent, traumatic and expensive ailments in human health care. The need for substitutes to replace or repair tissues or organs because of disease, trauma, or congenital problems is overwhelming. For example, in the United States alone, as many as twenty million patients per year suffer from various organ and tissue related maladies including burns, skin ulcers, diabetes, bone, cartilage, and connective tissue defects and diseases, more than eight million surgical procedures are performed annually to treat these cases, over 70,000 people are on transplant waiting lists, and an additional 100,000 patients die without even qualifying for the waiting list. The financial cost to care for these patients has been estimated to as much as $400 billion annually on patient suffering from organ failure or tissue loss [1,2].

The financial cost to care for these patients has been estimated to as much as $400 billion annually on patient suffering from organ failure or tissue loss [1,2]. To address this issue, Tissue Engineering, integrating a variety of742Submitted to: SFF Austin 2004 By Dr Wei Sunscience and engineering disciplines seeks to create functional tissues and organs for transplantation which restore, maintain or improve the function of human tissues, and evolves as one of the most promising therapies in the regenerative medicine [3].In the success of tissue engineering, three-dimensional (3D) tissue scaffolds play a critical role as extra-cellular matrices onto which cells can attach, grow, and form new tissues.

In the growth and migration processes that give rise to tissue function and morphogenesis, cells are influenced by a wide variety of factors including their own genetic programs, its bio-molecular composition, and 3-dimensional scaffolds to direct and induce cell proliferation, migration and differentiation. Modeling, design and fabrication of tissue scaffolds to meet multiple biological and biophysical requirements is always a challenge in tissue engineering. This is further amplified when designing load bearing scaffolds for bone and cartilage tissue application. In these cases, tissue scaffolds need to be designed with intricate architecture, porosity, pore size and shape, and interconnectivity in order to provide the needed structural strength, transport nutrients, and the micro-environment for cell and tissue ingrowth [4-6].

Thus, the designed scaffolds are often required to be made by advanced manufacturing process, such as solid freeform fabrication (SFF) due to the design complexity [7-9].The scaffold micro-architecture is believed to influence the behavior of cells and the biological function of tissues by providing a nutritional pathway as well as a spatial distribution for cell growth [1, 8, 9]. Designing such micro-architectures within scaffolds has always been limited due to limitations of current CAD technologies either in representation of such micro- structures or the inability to transfer microstructure information to RP machines. The objective of this paper is to develop a process tool path for freeform fabrication of biomimetic designed tissue scaffolds, including a novel Internal Architecture Design (IAD) approach which enables the generation of internal scaffold pattern from characteristic architecture patterns of the scaffold building blocks.

The output of such a process would be to have scaffolds with intended interior architecture which would ensure the right pore size, pore shape and interconnectivity throughout the scaffold while making use of current CAD technologies. The next section details current scaffold design techniques and its limitations and the motivation for the development of the IAD technique. The third section details the IAD approach giving details regarding the steps involved. The fourth and fifth section details the implementation of the algorithm and the sample models that were fabricated using this technique. For demonstration purposes we have used the 3DP™ based TheriForm™ Rapid prototype machine in fabrication of the model scaffolds.2.

For demonstration purposes we have used the 3DP™ based TheriForm™ Rapid prototype machine in fabrication of the model scaffolds.2. Current Scaffold Design Process and its LimitationsWith the advancement of SFF to be the most favorable approach in the fabrication of scaffolds for tissue engineering applications, CAD has been the default platform in which scaffolds were designed and then converted to machine instructions for fabrication. STL format which has been the de-facto standard of CAD model transfer to RP machines has been mostly used during the scaffold design and fabrication process. The interior architecture of these scaffolds were either designed as a pattern of extrusions, cuts and holes across the surface of the scaffolds in a CAD platform or its interior slice layers were filled in with contours or roads to generate the interior architecture.

The former method used by 3DP, SLS, SLA based RP machines for scaffold fabrication while the latter method is used by FDM based machines. Fig743Submitted to: SFF Austin 2004 By Dr Wei Sun1(a) and (b) outlines the method by which scaffolds are designed in a CAD platform in published literature. Chu et. Al [11] has used an assembly of designed unit cells and then intersected with the scaffold to obtain the interior architecture. Nam et al have used a similar approach in defining the interior architecture. These methods use STL as a means of data transfer to RP machine for final fabrication. This method has limitations as outlined below:1. STL tessellation involving approximation of surfaces with triangular facets is incapable of representing the scaffold’s micro-level pore feature sizes and intricate internal architectures. For example, as model precision demands become more stringent, the number of facets required to adequately approximate the model will increase.

For example, as model precision demands become more stringent, the number of facets required to adequately approximate the model will increase. This usually ends up with a computing-inefficient STL file due to the extremely large file sizes and inefficient to be used for tool path generation and for freeform fabrication [12, 13];2. Database in the STL format cannot include other design intents within the model, for example, topology or internal material variations, and its spatial distribution etc.3.

Database in the STL format cannot include other design intents within the model, for example, topology or internal material variations, and its spatial distribution etc.3. Expensive computing power required when the number of features and cuts in the scaffold become large enough to cause a memory overhead for the CAD software used.Fig 1(a) Chu et al[11]: Scaffold structure library and assembly.Fig 1(b) Boolean operation to cut out architecture within scaffold [10].Fig 1(c) How can we design the interior structure of the scaffold – Is it possible?Although we can use CAD models to represent an individually designed unit cell (such as one shown in Fig. 1(c), CAD is not generally capable of representing a unit cell assembly, for example, using different unit cells to fill in the femur structure as shown in Figure 1(c).

1(c), CAD is not generally capable of representing a unit cell assembly, for example, using different unit cells to fill in the femur structure as shown in Figure 1(c). Since each unit cell is biomimetically designed with complicated international architecture, the combination of them may result in an even intricate scaffold internal architecture which is beyond the capability of any current commercial available CAD systems [10]. In addition, the STL format which serves as the current industry standard for the data exchange between CAD and freeform fabrication can not be applied to such designed structures.744Submitted to: SFF Austin 2004 By Dr Wei SunTherefore, there is a strong demand for developing a new solution to transfer the information of biomimetic designed scaffolds to a process tool path of freeform fabrication techniques and in the design techniques used in the design of scaffolds.

We propose to develop a novel solution, an Internal Architecture Design (IAD) approach, to generate process planning instructions for freeform fabrication of biomimetic designed tissue scaffolds. The novelty of the IAD is that it allows the generation of the scaffold process planning without the creation of a unit cell filled 3D CAD scaffold model, thus avoiding the difficulty of using CAD to represent the assembly of the unit cell models. The IAD approach stems from the principle of the decomposition and material accumulation of layered manufacturing [14]. The approach uses the characteristic patterns to represent the individual unit cell 2D interior architectures, and assembles the individual characteristic pattern to form a 2D layered pattern for the scaffold. The processing tool path is generated for every layer based on the layered scaffold pattern, and is then applied to instruct freeform fabrication machines to build 3D scaffold layer by layer.

The processing tool path is generated for every layer based on the layered scaffold pattern, and is then applied to instruct freeform fabrication machines to build 3D scaffold layer by layer. In this way, we do not have to create a real 3D CAD based scaffold model, but still be able to design and control the internal architecture of the scaffold. Detailed development of the IAD approach is outlined as follows.3. Proposed Methodology for Interior Architecture Design of ScaffoldsStep 1: Determine the layered processing plane (zyxϕ , ),()According to the ordered sequences, the 3D volumetric scaffold V is sliced into layers during the model decomposition, and stacked back with materials during the material accumulation. The sequence is a number dispersal in 3-dimensional space V in which the layered manufacturing process is realized.

The sequence is a number dispersal in 3-dimensional space V in which the layered manufacturing process is realized. Let’s introduce a finite discontinuous real number set C to indicate the discrete layers which V is decomposed:C=⎧ ⎪ a;c ⎨ k ⎪⎩≤c1<c2<L<c1k −<ck<c1k +<L<cn<c1n +≤,2,1k,b =Ln,⎫ ⎪ ⎬ ⎪⎭(1)where subscript k represents the number of the kth decomposition in total n layers, and constants a and b represent the values of lowermost and uppermost sequence for a particular object V.Therefore, corresponding to each Ck, we can define a sequence function kφ (x, y, z), and assume it is equal to the ordered number ck,. as shown in Figure 2, we thus define a series of planes in V:ϕk,( zyx ,)=ck,k=L,2,1n(2)We call those planes the layered processing planes and represent them by a set of scale function kφ (x, y, z), with k = 1, 2, …n..

It is important to note that the layered processing plane kφ (x, y, z) represents both the indication of the processing layer, and the layered exterior geometry represented by CAD model for the scaffold. In the model decomposition process, the layered processing planes intersect and slice the designed 3D scaffold. In the material accumulation process, the layered processing planes form the processing layers on which material will be added. For example, in Therics’ TheriForm fabrication process [15-17], the745Submitted to: SFF Austin 2004 By Dr Wei Sunlayered processing plane is the plane on which the fresh biomaterial powder is spread and printed, while in Drexel’s multi-nozzle biopolymer deposition process, the layered processing plane is the plane on which the biopolymer is extruded and deposited.

The concept of the decomposition and accumulation, and the layered processing plane is further depicted in Figure 4.kφ (x, y, z) in the volumetric scaffold, we define Sk as a 2D layered scaffold pattern, which can be considered as unions of the partitioned unit cell characteristic patterns: k SOn each thus defined layered processing planek S 1=k 2k 3k iSSS(3)UUULUFor simplicity, we assume that the thickness of the slicing layer for unit cells and for thescaffold are all uniform and the same.Designed Scaffold ModelDecomposing (Slicing)Procedure of Data ProcessingModel DecompositionLayered Process InformationLayered Manufacturing PathLayered ManufacturingProcedure of Physical RealizationMaterial AccumulationLayer StackingFabricated ScaffoldFigure 2: Decomposition and accumulation for layered manufacturingThe external shape of the scaffold is defined by Non-Uniform Rational B-Spline (NURBS) surfaces.We use a direct slicing method to determine the intersection curves of the NURBS model with the given layer (k).

Figure 3 shows a step by step procedure on how to calculate the NURBS intersection points. As shown in the figure, a given NURBS surface (depicted as a 2D NURBS curve) is geometrically refined as an adaptive subdivision represented by many smaller domains. We apply bounding boxes to cover up these smaller domains, and find the box that intersects with ray which cast out on the slicing plane. A root finding procedure is initiated to converge at the intersection point. The procedure is then repeated for all rays that are cast onto the slice plane and for every slice plane that intersects the NURBS model.

The procedure is then repeated for all rays that are cast onto the slice plane and for every slice plane that intersects the NURBS model. By connecting the intersection point, we can determine the layered processing plane ( ) and hence determine the outer scaffold shape.zyxkϕ ,( ),746Submitted to: SFF Austin 2004 By Dr Wei SunFigure 3: Step by Step procedure on calculation of NURBS intersection pointsStep 2: Generate Toolpath based on Internal Architecture Design (IAD)The layered processing toolpath for freeform fabrication is then generated based on the 2D layered scaffold pattern. The above process can be briefly summarized into the following major steps:1. Define sub-volume Vi (unit cell) and the spatial position Pi; the discrete layers of V, for scaffold exterior zyxkϕ ,( ),2. Determine the unit cell characteristic patterns jiS ( i = 1,2, …m, and j =1,2, …) by slicingthem;3.

Determine the unit cell characteristic patterns jiS ( i = 1,2, …m, and j =1,2, …) by slicingthem;3. Form a 2D layered scaffold pattern kS(k=1, 2, …n) (or layered Interior Pattern) for agiven scaffold sliced layer by a union of all the unit cell patterns;4. Perform a intersection Boolean operation k4. Perform a intersection Boolean operation kbetween the 2D layered scaffold pattern and the exterior scaffold geometry given by φ to remove the unwanted regions of the scaffold sliced layer;(φ∩, zyx5. Conversion of the Interior Scaffold Pattern to process tool path information for freeform fabrication.6. Repeating Step 3 to Step 5 till k = n.We use a square block type unit cells with open pore in the center (as shown in Figure 4) as examples to illustrate the above process. Let’s assume that the designed scaffold consists of m unit cells with 4 unit cells being assigned to the top of the scaffold volume (Figure 4).

Let’s assume that the designed scaffold consists of m unit cells with 4 unit cells being assigned to the top of the scaffold volume (Figure 4). Assume that the number of slices of the scaffold has been determined, the layered processing planes (k = 1, 2, …n) for the scaffold exterior geometry are known, and the characteristic zyxkϕ ,( ), patterns of the unit cells have been computed and available in the database. The unit cell id Pi within different regions on this layer is used as a key to retrieve the appropriate unit cell characteristic patterns. A scenario has been illustrated as shownjiS corresponding to thezyxkϕ ,( ),747Submitted to: SFF Austin 2004 By Dr Wei Sunin Figure 4 by taking the case of 2 scaffold slice layers shown as slice layer 1 and slice layer 2.

2S The 2D layered scaffold patterns for scaffold layer 1, , are generated by 2 1 iS (i = 1, 2, … 4), respectively and are iS and union four unit cell characteristic patterns shown in the right side of Figure 5. Printing maps consisted of the geometric raster pattern for 2S , as shown in the figure. the slice layer 1 and slice layer 2 are then generated based on 1S and1S , and layer 2,By applying intersection Boolean operation between the Interior Scaffold Pattern and the layered processing plane , we then obtain the layered manufacturing maps which will be used for the process tool path generation. The tool path can be generated based on the particular SFF process used. The proposed algorithms is currently implemented in the 3DP™ based SFF technology with possible extension into contour based multi nozzle deposition system.

The proposed algorithms is currently implemented in the 3DP™ based SFF technology with possible extension into contour based multi nozzle deposition system. The interior scaffold pattern is used as the initial starting point to generate a raster pattern by determining the entry and exit points for the print head in the case of the 3DP based machine. The calculated intersection points will serve as the starting point for the tool path information for a 3DP based SFF machine. For vector contour based machines, the same scaffold pattern will be used to generate contour vectors by joining the intersection points in a linear technique and forming closed loop vector tool paths.zyxkϕ ,( ),Figure 4: Methodology for Internal Architecture Design4. ImplementationThe steps and algorithms outlined above are implemented within a fabrication planning software to control the different parameters of the fabrication process.

ImplementationThe steps and algorithms outlined above are implemented within a fabrication planning software to control the different parameters of the fabrication process. The algorithms are implemented in Microsoft's .NET development environment using C# as the programming language. Figure 5 gives the different component modules with the functional flow of data within the fabrication preprocessing environment.748Submitted to: SFF Austin 2004 By Dr Wei SunSlicing Module: Once the STEP file of the Scaffold CAD model is read, its geometry and topology will be maintained by the 3D Kernel by which the NURBS surfaces of the model with their topology information will be extracted for the slicing process. Direct slicing of the NURBS surfaces is performed by a Slicing Module, of which the results are packaged into a Slicing Object. The B-rep features of the model are viewed using graphic kernels for easier manipulation and display.

The B-rep features of the model are viewed using graphic kernels for easier manipulation and display. Tools are provided for viewing the model and verifying slice layer information for errors that may result in erroneous fabrication. Each cross sectional layer is then extracted from the model based on the slicing parameters and then converted to machine job instructions. A print job database is also maintained for database records and future retrievals.Figure 5: Fabrication Planning ComponentsUnit Cell Pattern Generation Module: The RP-ml file (containing tool path) from the previous module acts as the input file in this stage. Each layer is extracted from the file along with the microstructure cell ID/IDs it is associated with. The unit cell corresponding to the unit cell ID is retrieved from the design repository and is sliced to obtain the 2-D unit cell pattern.

The unit cell corresponding to the unit cell ID is retrieved from the design repository and is sliced to obtain the 2-D unit cell pattern. This unit cell pattern is then layered out onto the retrieved layer through a pattern generation algorithm as explained in Figure 5. Once this is completed, the new pattern is intersected with the original pattern of the CAD model retrieved from the RP-ml file to obtain a new set of pattern raster/contour lines. These new set of patterns can be stored in a new XML file format that can be distributed across the web for other applications or converted to machine job instructions to run an SFF based RP machine.5.

Results749Submitted to: SFF Austin 2004 By Dr Wei SunUnit CellBlockScaffold- Layer PatternsSize: 5mm x 5mm x 5mmPore Size: 2.5mm x 2.5mm x 2.5mmFig 6(a): Square pore architecture assigned to a Square scaffold structureSize: 5mm x 5mm x 5mm Pore Size: 2.5mm cylinder holesFig 6(b): Circular Pore architecture assigned to a circular shaped scaffoldSize: 5mm x 5mm x 5mmPore Size: 2.5mm x 2.5mm x 2.5mmFig 6(c): A circular pore architecture assigned to an irregular shaped scaffoldPorosity50787842Fig 6(d): Sliced layers of a multi-unit cell scaffold structure750Submitted to: SFF Austin 2004 By Dr Wei SunThe algorithm was tested with a couple of simple to complex models as shown in Figures 6. The size of the selected unit cell was arbitrarily selected to be a 5mm cube which had characteristic pores of a square pore and cylindrical hole within them.

The size of the selected unit cell was arbitrarily selected to be a 5mm cube which had characteristic pores of a square pore and cylindrical hole within them. The blocks define the outer shape of the scaffold while the unit cells define the interior architecture of the scaffold. Both the blocks and the unit cells were sliced at equal intervals and stored. The algorithm does not require any huge memory requirements and its run time only depends on the number of slices that make up the block. The size of the unit cell can go further smaller and its size is selected depending on other factors such as the desired pore size, the capability of the selected RP machine and material used. Fig 6(d) shows the bone block structure filled with four different unit cell architectures. Fig 7(a) and (b) shows a cylindrical shaped and a bone scaffold designed using the IAD methodology and fabricated using the TheriForm™ machine.

Fig 7(a) and (b) shows a cylindrical shaped and a bone scaffold designed using the IAD methodology and fabricated using the TheriForm™ machine. The scaffolds were made out of alumina with a slice thickness of 0.48mm and sintering temperature of 80C. As can be seen, the scaffolds have the required internal architecture as defined by the unit cell. Redefining the unit cell model architecture results in a different interior structure for the scaffold models.Fig 7(a) Cylindrical ModelFig 7(b) Bone shaped model6. ConclusionThe developed IAD methodology provides a more advanced design of scaffolds since each layer can individually be designed to have the desired layer pattern. This pattern is retrieved from the selected unit cell and at the end of the process, the desired scaffold with the outer shape of the block and selected interior unit cell architecture is fabricated.

This pattern is retrieved from the selected unit cell and at the end of the process, the desired scaffold with the outer shape of the block and selected interior unit cell architecture is fabricated. The advantages of the process include:The developed approach can help in the design of Biomimetic scaffolds that by current methodologies are impossible to create using CAD or other software. The developed approach does not exclude the use of CAD but uses it as an aid in the design of such scaffolds with macro and micro architecture. Interior architecture control of scaffolds along with the provision of multiple material specification and multiple architecture can be specified;Direct transfer of CAD model data to RP machines avoiding the bottleneck inherent in the use of the STL format;751Submitted to: SFF Austin 2004 By Dr Wei SunThe design of outer scaffold shape and the design of its interior are separated and hence do not present a memory overhead for CAD software.

By separating their designs, new architectures with varying properties can be designed and hence aid in the fabrication of better scaffolds for tissue engineering applications.The layer patterns that are obtained do not work in the same way as laying down hatch patterns within the contour layers. These patterns come with a reason and are generated from the unit cells. The unit cell that needs to be assigned within the block structure is selected based on factors such as the required mechanical and biological properties that have been pre-defined. The process assumes that this selection process has been carried out prior to the generation of process planning instructions. Further details into how these unit cells are selected is described in Sun et al [18-19] and the readers are referred to them for further details.The described process can be further extended to generate process planning instruction for contour based RP machines such as the FDM and PED [20].

Although the library of unit cells that can be fabricated with these machines are limited, the authors believe that by proper selection of varied contour patterns and sizes, the desired scaffold properties can be obtained. The IAD process developed for the TheriForm machine will be further extended to include contour based machines to be able to fabricate scaffolds with new biopolymer materials and in micron ranges.7. AcknowledgmentsThe authors would like to acknowledge the research funding support by Therics Inc,Princeton, New Jersey.8. References1. Langer R. and Vacanti, J. P., “Tissue Engineering”, Science, 1993, 260, 920-926. 2. U.S. Scientific Registry for Organ Transplantation and the Organ Procurement and Transplant Network. Annual Report. Richmond VA: UNOS, 1990.3. “Tissue Engineering”, 1995 Annual Report of the Whitaker Foundation: Tissue Engineering, 1995. http://www.whitaker.org/95_annual_report/tissue95.html4.

Annual Report. Richmond VA: UNOS, 1990.3. “Tissue Engineering”, 1995 Annual Report of the Whitaker Foundation: Tissue Engineering, 1995. http://www.whitaker.org/95_annual_report/tissue95.html4. Zeltinger J, Sherwood JK, Graham DA, Mueller R, Griffith LG., “Effect of pore size and void fraction on cellular adhesion, proliferation, and matrix deposition”, Tissue Engineering, 2001; 7, 557- 572..5. Zein I, Hutmacher DW, Tan KC, Teoh SH., “Fused deposition modeling of novel scaffold architectures for tissue engineering applications”, Biomaterials, 2002; 23:1169-1185.6. Hutmacher DW, Schantz T, Zein I, Ng KW, Teoh SH, Tan KC., “Mechanical properties and cell cultural response of polycaprolactone scaffolds designed and fabricated via fused deposition modeling”, Journal of Biomedical Materials Research, 2001; 55:203-216.7. Sun W, Lal P., “Recent development on computer aided tissue engineering – a review”, Computer Methods and Programs in Biomedicine, 2002; 67: 85-1038.

Sun W, Lal P., “Recent development on computer aided tissue engineering – a review”, Computer Methods and Programs in Biomedicine, 2002; 67: 85-1038. Yang, S.F., Du, Z.H., Leong, K.F., and Chua, C.K., “Review: The Design of Scaffolds fFor Use in Tissue Engineering. Part 1. Traditional Approaches”, Tissue Engineering”, 7(6), pp. 679-690, 2001. 9. Yang, S., Leong, K., Du, Z. and Chua, C., “The design of scaffolds for use in tissue engineering. Part2. Rapid prototyping techniques,” Tissue Engineering, 2002; 8 (1):1-11.752Submitted to: SFF Austin 2004 By Dr Wei Sun10. Nam J., Starly, B., Darling A., Sun W., “Computer-Aided Tissue Engineering for Modeling and Design of Novel Tissue Scaffold”, CAD'04 International Conference, Pattaya Beach, THAILAND, May 24-28, 2004.11. Chua, C.K., Leong, K.F., Cheah, C.M., Chua, S.W., “Development of a tissue engineering scaffold structure library for rapid prototyping. Part 1: Investigation and classification”, Int J.

OPOLYMER COMPOSITES WITH DAIRY PROTEIN FOR USE IN ADDITIVE MANUFACTURINGMaia Delwiche1 and John Obielodan1,*1Dept. of Mechanical Engineering, University of Wisconsin-Platteville, Platteville, WI 53818Corresponding author: obielodanj@uwplatt.eduAbstractAs the popularity and versatility of additive manufacturing grows, so does interest in developing new materials, including biopolymers. Casein is a protein found in dairy and historically, has been used in food applications, but its use as a biomaterial for engineering structures is less common. This study investigates the development of composite materials for additive manufacturing with casein as a biomaterial filler. To observe the effects of casein on material properties, vat photopolymerization-based and fused filament fabrication-based matrix materials were combined with different weight fractions of casein. Test samples were fabricated to evaluate tensile properties.

Test samples were fabricated to evaluate tensile properties. Test results show a maximum increase of 4% for FFF and 34% for SLA in the stiffness of the materials with casein compared to the neat matrix materials. However, the composite materials showed between 12% and 54% reductions in ductility, and marginal decreases in tensile strengths. The preliminary results indicate viability and prompt further investigation into casein-polymer composites for additive manufacturing.Keywords: Additive Manufacturing, Biopolymer, Dairy, Casein, PolymersIntroductionThe applications of additive manufacturing (AM) technologies are broad, ranging from prototyping to small batch manufacturing, medical implants and construction, and more uses continue to develop [1]. With technological developments and research initiatives, manufacturing materials continue to evolve as well.

With technological developments and research initiatives, manufacturing materials continue to evolve as well. Traditionally, polymer AM utilizes petroleum-based materials which carry associated hazards to human health and the environment [2]. Examples of these materials include Acrylonitrile Butadiene Styrene (ABS), Polycarbonate (PC), Nylon, or UV-Curing Resins [2-4]. A few biobased AM materials like polylactic acid (PLA) [2], Polyhydroxyalkanoates (PHAs) [5], and epoxidized sucrose soyate (ESS) resins [6] exist but further development of affordable materials sourced from sustainable resources is important to the future of additive manufacturing [2, 4].Many composite polymers for AM based on renewable resources have been developed in recent past. Some use biomaterials as fillers in petroleum-based polymers and resins [1,7-9], and others as fillers in biopolymer matrices, such as PLA [10-15].

Some use biomaterials as fillers in petroleum-based polymers and resins [1,7-9], and others as fillers in biopolymer matrices, such as PLA [10-15]. Lignin, a natural polymer sourced from wood, is a common additive in composite materials and has been recently studied as a filler1759in materials for fabrication using various AM technologies. [7-9, 13-14]. Cellulose nanofibers have also been incorporated into biobased polymers such as PLA as a beneficial reinforcer [11, 15]. Another study demonstrated that kenaf fibers, a natural plant fiber, in a PLA matrix improved both mechanical and thermo-mechanical properties of the material [12]. A study examining casein- functionalized cellulose fibers in PLA found whole milk casein protein to be an effective dispersant and binder in the composite [15].This research is an exploratory study on the viability of incorporating casein protein filler into different material matrices for fabricating engineering structures using AM processes.

Casein is a protein that is abundant in dairy; it makes up about 80% of the protein found in cow’s milk [16]. Its applications are primarily food or pharmaceutical based, but it has historically been used for polymer production as well [17]. Casein molecules consist of hydrophobic and hydrophilic blocks which are able to interact with other molecules [15]. These blocks participate in the crosslinking of casein during the manufacturing processes [17]. In the early 20th century, to create casein was combined with plasticizer, formable bioplastics [18]. However, the procedure was very time consuming, and the products fabricated from the casein plastic were predominantly small items such as buttons and buckles [18]. More recent products developed using casein include adhesives, foams, gels, and some bioplastics [16].formaldehyde, or other chemicalsThe source of the casein filler is of long-term interest as well.

Significant volumes of wastewater are generated in the dairy industry, with cheese production, whey manufacturing, and milk processing being major contributors [19]. These wastewaters, from washing equipment or whey by-products, have casein contents that can be valorized [19]. Another source of casein that can be valorized is discarded milk from deliberate dumping of milk shipments by farmers or cooperatives [20-22] It is routine for small amounts of milk to be dumped due to seasonal overproduction and other factors such as weather or batch contamination [22]. However, large- scale dumping occurs when there is a major disruption to the dairy industry such as supply-chain issues or a dramatic decrease in market demand [20-21]. All cases of dumping result in economic losses to farmers and the nutrient-rich milk can produce negative effects on the environment [16, 21].

All cases of dumping result in economic losses to farmers and the nutrient-rich milk can produce negative effects on the environment [16, 21]. Casein protein can be extracted from the wastewater or discarded milk using precipitation or microfiltration and dried into powder [23]. Casein is used for some medical applications and human or animal consumption but often the wastewater contains contaminants and would require purification to reach medical or food-grade industry standards; however, for engineering-grade applications, the purity of the casein is less relevant [16]. Extracting the casein and producing bioplastics for AM would valorize the dairy industry wastewater, and create an alternate market for discarded milk, reducing financial loss and environmental impact.

The materials generated would be sustainable, and, by sourcing the casein from waste, would also be affordable and not compete with the demands of the food industry.In this work, two different AM technologies are examined: stereolithography (SLA) and fused filament fabrication (FFF). SLA printing fabricates parts using a UV-photocurable resins. Ultraviolet light is employed to locally cure the resin in thin layers to build three-dimensional solid objects. Post-processing or cleaning is often required following SLA printing to finish the parts.1760The material for FFF is a solid filament made of thermoplastic polymer.

Post-processing or cleaning is often required following SLA printing to finish the parts.1760The material for FFF is a solid filament made of thermoplastic polymer. The filament is fed into the print head and heated while being extruded through a nozzle and deposited in layers onto the build platform to form a part.The purpose of this research is to develop sustainable materials for AM and to reduce economic loss within the dairy industry by determining whether casein protein is a viable filler in composites development for AM. Casein powder will be added to neat polymer materials and test specimens printed. The mechanical properties of the composites will be compared to the neat polymer to observe the effects of the casein filler.MethodologyFor both SLA and FFF printing methods, the casein powder was blended with the neat matrix material in different weight fractions.

Five dog-bone tensile test specimens were produced for each treatment to evaluate the mechanical properties. Specimens were dried in a vacuum oven and stored in a desiccator prior to testing to prevent inconsistency in material behavior due to moisture absorption from the air. Tensile testing parameters followed ASTM D638 standards for plastics. A 50 kN MTS tensile testing machine was used for testing all specimens. After testing, the fractured surfaces of the FFF specimens were examined under a scanning electron microscope (SEM). Backscattered-Electron (BSE) imaging was used with variable pressure settings to examine the morphology of the fractured surfaces.SLA PrintingFor SLA printing, Formlabs Clear Resin was used as a matrix for composites containing 5 wt% and 10 wt% casein filler material. Tensile test specimens were printed on a Form 3 printer from Formlabs using the two resin-casein composites. The specimens were angled off the bed of the printer as shown in Fig.

Tensile test specimens were printed on a Form 3 printer from Formlabs using the two resin-casein composites. The specimens were angled off the bed of the printer as shown in Fig. 1 for ease of removal after print.Figure 1: SLA tensile specimens before removal from the printer bed.After printing, the samples were post-processed in a Formwash chamber with isopropyl alcohol to clean the specimens, then cured in a UV Formcure oven. Sample batches were cured at 60 oC for different times. The control samples made of neat, standard clear resin were cured for 151761and 30 minutes as recommended by Formlabs [24]. The 30 minutes cure was to achieve a marginal gain in mechanical properties.

The 30 minutes cure was to achieve a marginal gain in mechanical properties. The composite resin batch with 5 wt% casein was cured for 30 minutes while two batches containing 10 wt% casein were cured, one for 30 minutes and the other, for 60 minutes, to compare the effect of longer cure time on the properties of the composite.FFF PrintingPLA powder was blended with 5 wt%, 10 wt%, and 12 wt% casein powder and then extruded into filament for FFF printing. A lab-scale Filabot EX2 extruder was used for filament extrusion. The first extrusion was pelletized and re-extruded a second time to ensure the homogeneity of the final filament intended for printing. The pelletized filament can be seen in Fig. 2 and a sample of PLA/Casein filament is shown in Fig. 3. Tensile test specimens were fabricated on a Creality 10-S 3D printer. The specimens were oriented flat on the print bed, with the long edge parallel to the printer’s y-axis.

3. Tensile test specimens were fabricated on a Creality 10-S 3D printer. The specimens were oriented flat on the print bed, with the long edge parallel to the printer’s y-axis. They were printed with 100% infill settings and a zig-zag deposition pattern.Figure 2: pelletized PLA/casein filament after the first extrusion.Figure 3: PLA/casein filament sample.Results and DiscussionSLA SpecimensThe comparison plots for SLA printing are seen in Fig. 4 to Fig. 6 for the tensile strengths, moduli of elasticity, and percent elongations respectively. The designations below each bar on the chart indicate the composition of the material (CL for clear resin), followed by the weight percentage of casein added and the cure time in minutes, thus, CL05C30 indicates clear resin/5 wt% casein cured for 30 minutes. For neat clear resin samples, an increase from 15 to 30 minutes in cure time resulted in better performance across all three material properties measured.

For neat clear resin samples, an increase from 15 to 30 minutes in cure time resulted in better performance across all three material properties measured. Interestingly, with the addition of 10 wt% casein, the change in material properties was varied when comparing 60 minutes cure time to 30. The modulus experienced a 15% reduction while the elongation increased 27%, and the tensile strength remained constant.1762Samples with 0 wt%, 5 wt%, and 10 wt% casein were compared at a consistent cure time of 30 minutes. The tensile strength decreased by just under 15% for the composites compared to the neat resin, and the change was close to uniform for both weight fractions. The modulus of elasticity showed an increase with the addition of casein, 34% at 5 wt% casein and 13% for 10 wt% casein.

The modulus of elasticity showed an increase with the addition of casein, 34% at 5 wt% casein and 13% for 10 wt% casein. The percent elongation was reduced to nearly half of that of the neat resin with the reduction more marked for 5 wt% casein.Tensile Strengthsa P M70 60 50 40 30 20 10 0CL00C15CL00C30CL05C30CL10C30CL10C60MaterialsFigure 4: Tensile strength comparison for SLA samples.Modulus of Elasticity54a P G3210CL00C15CL00C30CL05C30CL10C30CL10C60MaterialsFigure 5: Modulus of elasticity comparison for SLA samples.1763%Elongation)m m m m/( n o i t a g n o E %l7 6 5 4 3 2 1 0CL00C15CL00C30CL05C30CL10C30CL10C60MaterialsFigure 6: Percent elongation comparison for SLA samples.FFF SpecimensFigures 7 to 9 show the comparison of the tensile strengths, moduli of elasticity, and percent elongations respectively for FFF specimens.

The five-specimen averages for PLA/casein composites made of 5 wt%, 10 wt%, and 12 wt% casein are shown in the plots. Samples of neat PLA were also tested as a control. A slight increase in the modulus of about 4% with the addition of 5 wt% casein was observed and overall, there was only a marginal change in the modulus across all PLA/casein compositions. The tensile strength and percent elongation both experienced varying levels of reduction. The decrease in percent elongation was more significant and ranged from 12% to 35%.

The tensile strength and percent elongation both experienced varying levels of reduction. The decrease in percent elongation was more significant and ranged from 12% to 35%. This decrease indicated that the addition of casein created a more brittle material compared to the neat PLA samples.Tensile Strength] a P M[60 50 40 30 20 10 0Neat PLA 2nd Extrusion5%10%12%wt% CaseinFigure 7: Tensile strength comparison for FFF samples.1764Modulus of Elasticity54] a P G[3210Neat PLA 2nd Extrusion5%10%12%wt% CaseinFigure 8: Modulus of elasticity comparison for FFF samples.% Elongation]m m m m/[3.5 3 2.5 2 1.5 1 0.5 0Neat PLA 2nd Extrusion5%10%12%wt% CaseinFigure 9: Percent elongation comparison for FFF samples.SEM MicrographsFigure 10 presents the SEM images of the fractured surface of FFF samples at several magnifications. Comparing the image of the neat PLA surface in Fig. 10a to the images for 5 wt% in Fig.

Printability was viable for FFF and SLA printing in the 5 wt%-12 wt% casein filler range examined. Initial testing showed varied mechanical property changes but overall, the composites were more brittle while demonstrating only a slight change to strength and stiffness. For both printing technologies, the 5 wt% casein composites performed the best. The printed material appears homogeneous in SEM imaging. Further refinement of the material blending and printing would improve the process. Additional study includes testing and examining of other weight fractions of casein and incorporating other components such as a plasticizer to reduce the brittleness of the materials.AcknowledgmentsThe authors gratefully appreciate the contributions of Jonas Gilles for his contribution in sample preparations, experimental setup, and data generation. This research was supported by Dairy Industry – Impact and Innovation - Faculty Fellowships through Dairy Innovation Hub (DIH).

This research was supported by Dairy Industry – Impact and Innovation - Faculty Fellowships through Dairy Innovation Hub (DIH). The views expressed herein are those of the authors and are not necessarily those of DIH.1766References[1]T. Ngo, A. Kashani, G. Imbalzano, K. Nguyen and D. Hui, "Additive manufacturing (3D printing): A review of materials, methods, applications and challenges," Composites Part B: Engineering, vol. 143, pp. 172-196, 2018.[2] M. D. Tabone, J. J. Cregg, E. J. Beckman and A. E. Landi, "Sustainability Metrics: LifeCycle Assessment and Green Design in Polymers," Environmental Science & Technology, vol. 44, no. 21, p. 8264–8269, 2010.[3] B. Wang, Z. Zhang, Z. Q. J. Pei and S. Wang, "Current Progress on the 3D Printing of Thermosets," Advanced Composites and Hybrid Materials, vol. 3, p. 462–472, 2020.[4]S. S. Alghamdi, S. John, N. R. Choudhury and N. K.

2, no. 3, p. 171–189, 2009.[18] G. H. Brother, "Casein Plastics," Industrial and Engineering Chemistry, pp. 31-33, 1940.[19] C. Costa, N. Azoia, L. Coelho, R. Freixo, P. Batista and M. Pintado, "Proteins Derived from the Dairy Losses and By-Products as Raw Materials for Non-Food Applications," Foods, vol. 10, no. 35, 2021.[20] C. A. Wolf, A. M. Novakovic and M. W. Stephenson, "COVID-19 and the U.S. DairySupply Chain," Choices, vol. 36, no. 3, pp. 1-8, 2021.[21] A. Weersink, M. v. Massow and B. McDougall, "Economic thoughts on the potentialimplications of COVID-19on the Canadian dairy and poultry sectors," Canadian Journal of Agricultural Economics, vol. 68, no. 2, pp. 195-200, 2020.[22] A. M. Novakovic and C. A. Wolf, "Disorderly Marketing in the Twenty-First CenturyU.S. Dairy Industry," Choices, vol. 33, no. 4, pp. 1-6, 2018.[23] B. Carter, N. Cheng, R. Kapoor, G. Meletharayil and M.

Novakovic and C. A. Wolf, "Disorderly Marketing in the Twenty-First CenturyU.S. Dairy Industry," Choices, vol. 33, no. 4, pp. 1-6, 2018.[23] B. Carter, N. Cheng, R. Kapoor, G. Meletharayil and M. Drake, "Microfiltration-DerivedCasein and Whey Proteins From Milk," Journal of Dairy Science, vol. 104, no. 3, pp. 2465-2479, 2021.[24]"Form Cure time and temperature settings," Formlabs, 2 March 2021. [Online]. Available: https://support.formlabs.com/s/article/Form-Cure-Time-and-Temperature- Settings?language=en_US#settings. [Accessed 16 August 2022].1768.

oresorbable Implants using Selective Laser Melting S. Hoeges*, M. Lindner†, W. Meiners* and R. Smeets§Fraunhofer Institute for Laser technology, Aachen, Germany† Dental Materials and Biomaterials Research, University Hospital Aachen, Germany§ Department of Oral and Maxillofacial Surgery, University Hospital Aachen, GermanyAbstract Using bioresorbable materials implants can be manufactured which dissolve in the human body and are replaced by natural bone structure. For large implants an interconnecting porous structure needs to be integrated in the implant for a good vascularisation. Using additive manufacturing technology these internal structures can be directly manufactured. The structure can be designed by consequent following the guidelines of the medical expert. This paper describes the develop- ment of Selective Laser Melting to process bioresorbable materials Poly(D,L-lactide) and - Tricalciumphosphate.

This paper describes the develop- ment of Selective Laser Melting to process bioresorbable materials Poly(D,L-lactide) and - Tricalciumphosphate. The properties of the parts concerning microstructure, mechanical and biological properties after processing are analyzed in laboratory and animal tests. Possible applications are demonstrated and include individual bone substitute implants in cranio-maxillo- facial surgery.IntroductionIn medical science rapid prototyping (RP) and rapid manufacturing (RM) are manufacturing technologies with rapidly growing influence and large potential for several applications. Some applications are already industrial standard with growing numbers of applicants. These are for example the manufacturing of teeth restorations out of Cobalt-Chromium-alloy or the manufac- turing of complex orthopedic devices. The manufacturing of bone substitute implants using RM technology is another field of interest.

The manufacturing of bone substitute implants using RM technology is another field of interest. Additive Manufacturing like Selective Laser Melting (SLM) (also described as Direct Metal Laser Sintering or LaserCusing) or Electron Beam Melt- ing (EBM) has advantages to conventional manufacturing like milling or casting concerning geometrical freedom, flexible production, material consumption and manufacturing time. The manufacturing with additive technology becomes economical in two cases. The first case is the manufacturing of individual implants in small series. There is no need for the casting or deep drawing. Individual implants based on CAD-data can be directly manufactured without changes in machine set up. Secondly manufacturing becomes economical when new functionalities can be integrated in the implant which would not be possible by conventional manufacturing.

Secondly manufacturing becomes economical when new functionalities can be integrated in the implant which would not be possible by conventional manufacturing. When these functionalities result in better quality of life of the patient, higher production costs can be negotiated be better durability and functionality of the implant. These functionalities can be forexample controlled porosity of the implant, adapted mechanical properties or new composite materials.There have been many investigations on the manufacturing of permanent medical implants out of titanium(alloys), cobalt-chromium-alloys or implant steel [1, 2]. There were already clinical studies and applications of implants in the human body which were manufactured by Rapid Manufacturing technologies [3]. In many cases regenerative methods would be much better suited for the treatment of bone defects.

In many cases regenerative methods would be much better suited for the treatment of bone defects. There is taking place a strategical change in implanto- logy from permanent implants to regenerative methods using resorbable materials. The gold standard at the moment is human bone taken e.g. from the iliac crest. This method requires a second surgery with its secondary morbidity. To avoid this, bioresorbable bone substitute ma- terials are developed. There are several products on the market using different materials and different processing technologies. The materials can be used by the surgeon as paste, granules or semi-finished parts like cuboids or wedges. At the moment there is no processing technique to manufacture individual implants out of bioresorbable materials. This is based on two reasons. Conventional manufacturing methods involve a sintering process. This requires a tool to shape the implant.

This is based on two reasons. Conventional manufacturing methods involve a sintering process. This requires a tool to shape the implant. Individual implants would therefore be expensive due to expensive tool production. The second reason is the internal structure of bioresorbable implants. For full vascularisation of large implants an interconnected porous structure needs to be integrated in the implant. By varying the pore size ingrowth of cell tissue into the implant is possible. To realise a full vascu- larisation the pore size and interconnectivity needs to be well defined. These two problems for manufacturing of individual bioresorbable implants can be solved by rapid manufacturing tech- nologies.

These two problems for manufacturing of individual bioresorbable implants can be solved by rapid manufacturing tech- nologies. The geometrical freedom of additive manufacturing makes the production of individual geometries and the realisation of a defined interconnected porosity possible by integrating the porous structure in the design of the implant.There are several bioresorbable materials for example polymer-based [4], bioglass [5], resorb- able metals or bioceramics [6, 7]. Bioceramics based on calcium phosphate (e.g. -TCP) have the highest acceptance in medical science since their chemical structure is closest to human bone. In this work the additive manufacturing with Selective Laser Melting (SLM) is applied. For fusion of the powder a melt phase is required. Materials with thermal instability (like -TCP) cannot be processed with SLM directly. Materials suitable for processing using melting of the material are polymer based materials e.g. polylactide (thermoplastic).

Materials suitable for processing using melting of the material are polymer based materials e.g. polylactide (thermoplastic). During the resorption of polylactides an acidic environment is generated next to the implant due to the lactidic structure of the resorbable polymer. An acidic environment in human tissue results in inflammatory reactions [4].For optimised properties of manufactured implants a combination of both materials is chosen: polylactide for processing using SLM and -TCP for optimised biological properties. The approach of combining two materials has other medical advantages.

The approach of combining two materials has other medical advantages. The acidic surrounding can be decreased by using a composite with alcaline materials like calcium phosphate ceramics.Using a composite of polymer and ceramic material the mechanical properties of the parts can be enhanced to pure polymer [8].The main challenges for manufacturing bioresorbable implants using SLM are the preparation of the used materials for the SLM process and the processing of the material without damaging or changing its structure and chemical composition.This work deals with the process adaption of Selective Laser Melting to process the bioresorb- able composite poly(D,L-lactide) and completely new approach using medical approved materials to generate parts by full melting of PDLLA and embedding of -TCP in a matrix of PDLLA.tricalcium phosphate (PDLLA/TCP).

This is aMaterials and MethodsReflection of radiation The processing of the used material with SLM strongly depends on the optical properties of the material. The laser source (wavelength) is then used depending on the measured reflection spec- trum. As the material was used as a powder, the reflection of radiation of different wavelengths is analyzed on powder material. For wavelengths ranging from 400-2600 nm an UV/VIS/NIR Spectrometer Lambda 9 from Perkin Elmer is used. For wavelengths from 1900-22000 nm a FT- IR Spectrometer 1725 X from Perkin Elmer is used.Powder morphologyTCP powder (Biovision GmbH Biomaterial, Wiesbaden, Germany) and PDLLA (BoehringerIngelheim, Ingelheim, Germany) was used to produce the composite powder. Both materials were filled in a polyethylene bottle. Zirconia milling balls were added and the bottle was put on a rolling platform for three weeks. The resulting powder was prepared with a sieve of 90 µm mesh.

Zirconia milling balls were added and the bottle was put on a rolling platform for three weeks. The resulting powder was prepared with a sieve of 90 µm mesh. The prepared composite powder was characterized before and after the milling process. The particles size distribution was measured using a particle size analyzer (Mastersizer 2000, Malvern, Worcestershire, Great Britain). The milled particles were embedded in a resin and a cross-section was prepared and analyzed using scanning electron microscopy (LEO 440i, Carl Zeiss SMT AG, Oberkochen, Germany).SLM processing For the processing of the composite material PDLLA and -TCP using SLM, a process chamber developed at ILT was used together with a CO2-laser (FEHA 400S, FEHA, Germany). The laser beam was projected on the powder bed using a scanning device (hurryscan 20, Scanlab, Germany) and an F-Theta focusing lens. The CO2-laser operates in cw-mode at a wavelength of 10600 nm.

The laser beam was projected on the powder bed using a scanning device (hurryscan 20, Scanlab, Germany) and an F-Theta focusing lens. The CO2-laser operates in cw-mode at a wavelength of 10600 nm. To adjust the laser power for low intensities (0.2-10 W) a power attenuator (ULO optics, United Kingdom) was used. A schematic of the experimental set up is shown in Fig 1. To achieve a reproducible deposition of powder layers the powder deposition device was adopted.For qualifying the SLM process for the processing of the composite material an adaption of the process parameters laser power, laser beam diameter, scanning velocity, track distance as well as powder layer thickness was done.Fig 1: Schematic of the experimental setupMicrostructural analysis One problem of powder based additive manufacturing is the appearance of pores and defects in the microstructure of the generated parts. This results in poor mechanical properties.

This results in poor mechanical properties. Therefore the first aim for qualifying SLM to process the composite material PDLLA/ -TCP is to produce dense parts with a relative density >98%. Analysis of the density is done by optical analysis of cross sections of parts generated with different process parameters. An optical microscope and an analyzing software (Analysis, Olympus, Hamburg, Germany) is used to perform the measure- ment of the density.Mechanical properties The mechanical properties of bioresorbable bone substitute implants are tested with compression tests on porous test samples. This is in correlation to the literature [9, 10].

This is in correlation to the literature [9, 10]. Cubes with varying pore diameter (500, 600 and 800 µm) and resulting porosity are manufactured using SLM and compressed until fracture of the part.Biological analysis The implementation of new manufacturing technologies for the processing of medical approved materials for series production of implants requires biological and medical analysis of manufac- tured parts. In biological tests (in vitro tests) the effect of the generated parts on bone cells (osteoblasts) is analyzed. Bioresorbable test geometries will be manufactured with an inter- connecting pore structure to induce the ingrowth of bone into the implant for full resorption of the part. The pore diameter shows significant influence on the growth behavior of the cells. Test parts with different pore diameters are manufactured and the proliferation of the cells on the parts is measured using life / dead staining of the cells and optical analysis.

Test parts with different pore diameters are manufactured and the proliferation of the cells on the parts is measured using life / dead staining of the cells and optical analysis. To minimize the geometrical influence of the parts on the results, the same test geometry is used for in vitro andin vivo tests (animal tests). The geometry is a cylinder of 15.6 mm diameter and 5 mm thickness. Pore structure is designed using CAD software according to Fig 2 using multiplied unit cells. The pore diameter is varied between 500-800 µm.Fig 2: Schematic of the design of porous test geometries using multiple unit cells.The test geometries are manufactured using process parameters for manufacturing of dense parts without defects to achieve high strength of the parts. The powder remaining in the pore structure is then removed by blasting with alumina particals. The manufactured and blasted test parts are then cleaned in ethanol, -sterilized and populated with human osteoblasts (HOB).

The manufactured and blasted test parts are then cleaned in ethanol, -sterilized and populated with human osteoblasts (HOB). The analysis of the cultured cells takes place after 1, 3 and 6 weeks. The results are compared to reference material (poylvinylalcohol, PVA) and phase pure medical grade -TCP. The cells are then counted using optical analysis.Animal experimental studies Results concerning degradation and bone ingrowth into porous implants can only be achieved by suitable models of animal studies. To evaluate the properties of SLM implants with inter- connected porous structure in vivo 16 implants as described in the previous section with a pore channel diameter of 600 µm are manufactured. The parts are then sand blasted, washed in ethanol and -sterilized as described earlier. The animals are 20 rabbits of the race Chinchilla. The bone defect is located in the skull (calvarium) of the animals as shown in Fig 3.

The animals are 20 rabbits of the race Chinchilla. The bone defect is located in the skull (calvarium) of the animals as shown in Fig 3. Four defects are filled with autologous bone (control), 16 defects are provided with SLM implants.Fig 3: Schematic of the bone defect and the skin incisionHistological analysis is performed on cross sections of the test area after 8 weeks.ResultsReflection of radiation Fig 4 shows the measured diffuse reflection coefficient on composite powder out of PDLLA / - TCP. At =10600 nm (CO2-laser) a reflection of radiation of 4.5 % is measured. No transmission of radiation on powder layers of thickness > 30 µm is observed. Therefore 95.5 % of the radia- tion is absorbed in the powder layer and is used for the melting of the powder layer. The absorp- tion coefficient at wavelength =1064 (Nd:YAG-laser) is between 2-6 %.

Therefore 95.5 % of the radia- tion is absorbed in the powder layer and is used for the melting of the powder layer. The absorp- tion coefficient at wavelength =1064 (Nd:YAG-laser) is between 2-6 %. The process of SLM is therefore expected to be efficient and more stable using CO2-laser radiation.Fig 4: Reflexion of radiation at wavelength =2500 – 22500 nm for composite powder materialThe results were confirmed using the SLM process chamber with different laser sources (diode pumped Nd:YAG laser ( =1064 nm) and CO2-laser ( =10600 nm)). Much more laser power (tenfold) is needed to melt the polylactide using Nd:YAG-laser radiation. Once the material is molten, absorption increases and the material is burned. No suitable processing window could be identified to generate parts using Nd:YAG-laser radiation.Powder morphology Fig 5 shows SEM-pictures of the prepared composite particles out of PDLLA and -TCP. Particles show a homogenous size distribution and are of cuboidal shape.

Particles show a homogenous size distribution and are of cuboidal shape. To ensure a homo- genous distribution of the composite materials in the generated part, homogenous distribution of the materials in each particle is aimed for to prevent separation during powder layer deposition. By grinding both materials in one process, the smaller -TCP particles are ground into the PDLLA particles (Fig 5 b)).a)b)Fig 5: Composite particles out of PDLLA / -TCP a) SEM picture, b) cross sectionThe particle size distribution is shown in Fig 6. The first maximum of particle size results of - TCP particles which are between 1-5 µm of size. The second maximum at 35 µm is given by the PDLLA particles.Fig 6: Particle size distribution for composite material out of -TCP / PDLLAThe particle size distribution is suitable for the SLM process since the particles are smaller than one powder layer (50-100 µm).

The flowability of the particles is limited, a suitable powder deposition technique needs to be developed.Powder deposition Due to the grinding process the powder particles are not spherical with poor flowability compared to spherical metal particles which have been used for SLM in the past [1]. For a reproducible powder layer deposition the powder deposition device is varied. The powder layer is crucial for the manufacturing of dense parts. Each layer has to be dense and of a defined thick- ness varied between 30 – 100 µm.During SLM of metals the powder is deposited using a carbon fibre brush or a metal blade. This works well with spherical particles of high density (e.g. metals). Transferring these results to the deposition of the composite material a new deposition device is essential. The reason is the poor flowability of the particles and the low packing density due to low density of PDLLA (1.2 g/cm³).

The reason is the poor flowability of the particles and the low packing density due to low density of PDLLA (1.2 g/cm³). Therefore a powder deposition with additional compression of the powder layer needs to be developed. One attempt is the use of a bended aluminium plate (see Fig 7 a)). This approach results in a powder layer shown in Fig 8 a). The powder layer is of defined thickness, the com- pression is very high. Due to inhomogeneous compression of the powder bed powder ripples appear. This technique is not suitable for reproducible powder deposition. Different approaches with different angles of the plate surface to the powder bed resulted in no significant improve- ment in the appearance of ripples. Another approach is using a counter rotating cylinder. The schematic is shown in Fig 7 c).

%]0,2 - 105 - 50010 – 20020 - 100< 50030 - 70Tab 1: Variation of process parametersFig 9 a) shows an SEM image of a cross section of a test part. Single pores can still be observed, the overall density is >98 % in the processed area. This result is the basis for manufacturing of demonstrational implants. One application is an implant for a large bone defect in the skull with individually fitted design to reproduce the original shape of the bone. A demonstrational implant is manufactured using parameters leading to nearly full density of the part (Fig 9 b).a)b)Fig 9: a) Cross section of test geometry out of PDLLA / -TCP, b) Individual cranial implantMechanical properties The compression strength of the test geometries is shown in Fig 10. Depending on the porosity of the test parts compression strength between 3.4 – 8.8 MPa were achieved. As supposed, parts with higher porosity (larger pore diameter) show lower strength.

As supposed, parts with higher porosity (larger pore diameter) show lower strength. The strength of the parts is in the area of medium to high strength of cancellous bone (2-12 MPa [11]).Pore diameter dP [µm]Porosity P [%]Compression strengthkomp [MPa]0,5 0,6 0,835 47 538,80 ± 0,05 4,22 ± 0,51 3,43 ± 0,77Tab 2: Compression strength of porous test geometries with different pore structureSince the strength of the parts is comparable to cancellous bone SLM creates the possibility to manufacture bone substitute implants for non load-bearing bone defects which are suited for clinical use.Biological analysis (in vitro) The test geometries for in vitro and in vivo analysis manufactured with SLM are shown in Fig 10.5 mmFig 10: Manufactured parts with interconnecting pore structure for in vitro and in vivo analysisThe results of the in vitro analysis are shown in Fig 11.

The graph shows the relative fluores- cence of the living cells, greater fluorescence therefore indicates a higher number of living cells. The proliferation of osteoblasts exceeds for SLM-parts with pore size 600 µm and 700 µm the results achieved on the control and phase pure -TCP. Since no conclusion can be drawn out of this result between 600µm and 700 µm, 600 µm is chosen for animal testing which is in good accordance with the literature [2, 12]). These results show as well that the SLM process has no toxic influence on the proliferation of bone cell.Fig 11: Proliferation of osteoblasts on SLM test geometries for different pore sizesAnimal studies (rabbits) The removed implants are analyzed on the presence of new bone formation. Histological sections are shown in Fig 12. After 12 weeks natural bone minerals (NBM) can be seen in Fig 12 a).

Histological sections are shown in Fig 12. After 12 weeks natural bone minerals (NBM) can be seen in Fig 12 a). New bone formation can be observed in Fig 13 b) (arrows).NBMNBMa)b)Fig 12: Histological sections (200x) of implants showing natural bone minerals (a) and new bone formation (b)Histological analysis show, that new bone formation has started after 8 weeks with bone form- ation in critical size defects, i.e. bone defects that do not heal without scaffold. This demonstrates the potential of future application of SLM implants for healing large bone defects.DiscussionThe presented results show, that bioresorbable bone substitute implants can be manufactured with Selective Laser Melting (SLM). Particles out of a composite material of poly(D,L-lactide) (PDLLA) and -Tricalciumphosphate ( -TCP) have been prepared by grinding. Particles show homogenous distribution of the composites and are suitable for powder layer deposition during additive manufacturing.

Particles show homogenous distribution of the composites and are suitable for powder layer deposition during additive manufacturing. SLM process has been developed to generate dense parts with a homo- genous distribution of composites without major changes in physical and chemical properties of the materials. For full vascularisation of bioresorbable implants an interconnecting porous struc- ture is necessary. Using additive manufacturing technology SLM, pore structures can be inte- grated in the design of the implant and are reproduced in the part. To demonstrate the possible application of the manufacturing technology a skull defect is used to manufacture an individually designed implant with interconnecting porous structure shown in Fig 13.The mechanical properties of fabricated test geometries are in the medium to high range of human cancellous bone and the implants are therefore suitable for application in non load-bear- ing bone defects.

In vitro experiments show good proliferation and no cytotoxicity of human osteoblasts on SLM test geometries. This result demonstrates that no changes of the biological material properties occur during the processing of the material. Following these positive results animal studies have been carried out to analyze the new bone formation in the implant in vivo. Histological analysis shows that natural bone minerals and new bone formation takes place during 8 weeks in the scull of a chinchilla. Bone ingrowth is observed into the porous structure of the test geometries. These first results show the high potential of porous structures manu- factured by SLM for the reconstruction of large bone defects. Still more and large scale studies have to be carried to confirm the potential.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference28th Annual International Solid Freeform Fabrication SymposiumBlown Powder Laser Cladding with Novel Processing Parameters for Isotropic Material PropertiesJ. Liu1, E. Fearon1, 2, S.P. Edwardson1, G. Dearden11 Laser Engineering Group, Mechanical, Material & Aerospace Engineering, School of Engineering, the University of Liverpool, L69 3GH 2 Systems and Products Development, Advanced Laser Technology Ltd, Manchester, M1 2NPAbstractA limitation for blown powder laser additive manufacturing in many applications is the material properties of parts made tend to show anisotropy due to directional solidification of the microstructure. Recent work reported here has identified novel low power processing conditions that yield equiaxed grain structures in 316L stainless steel and thus potentially eliminate material anisotropy.

Recent work reported here has identified novel low power processing conditions that yield equiaxed grain structures in 316L stainless steel and thus potentially eliminate material anisotropy. Initial observations show that the process window is affected by build height, substrate choice, powder, powder delivery rates, laser power and processing speed. A system has been developed to give precision layer height control via nozzle design and low powder delivery rates through an in-house design of powder hopper. Mechanical tests have been conducted under the novel processing parameters. Large amounts and uniform distribution of equiaxed structures compared to standard process conditions in 316L are found significantly. Moreover, cladding has been successfully produced with significantly low power around 350W, thus potentially improving process efficiency and set-up cost1.

Moreover, cladding has been successfully produced with significantly low power around 350W, thus potentially improving process efficiency and set-up cost1. IntroductionAdditive manufacturing (AM) is receiving unprecedented attention from the mainstream media, the investment community, and national governments around the world. While there are multitude of laser based AM applications, laser cladding is the focus of this work. The Laser Cladding (LC) process can be known by many different names such as Laser Free Form Fabrication (LFFF), Direct Metal Deposition (DMD) or Directed Light Fabrication (DLF) (Mazumder, 2017).Figure 1: Apparatus of Laser Cladding at Liverpool with customised 4-port directional nozzles.There are two main families of additive manufacturing technologies for production parts, powder bed technologies and blown powder technologies. Blown powder laser cladding (BPLC) is mainly explored in this work.

Blown powder laser cladding (BPLC) is mainly explored in this work. In this process, metal powder is blown either coaxially or via a side feed into a laser generated melt pool, adding to its volume to form a fusion-bonded raised bead when cooled. This can then be built up layer by layer in a net shape process (Figure157228th Annual International Solid Freeform Fabrication Symposium1). This offers all the advantages of AM such as increased design freedom, light weight structures, new functionality and short production cycle times. In addition to powder bed technologies, BPLC also offers the addition of functionalities on existing parts with the same or a different material, compositional variation, repair of parts high value parts and no dimensional limits apart from machine size.The application of LC also has advantages in aerofoil manufacturing with lower weight but high strength to weight ratio.

This helps with shortening processing periods by around 80% as well as 20-50% reduction in cost (Ma, et al., 2013). Compared with conventional processing, laser cladding can provide high wear and corrosion resistance in terms of coating which is applied in the engine manufacture by Rolls-Royce and General Electric. In addition, rapid prototyping benefits from this technology.A recent study on blown powder laser cladding by Song. et. al (Song, et al., 2016) has proved that addition of graded structures on existing part of different materials can clad. They applied Ni50Cr as blown powder cladding on 304 stainless steel successfully. In terms of different exploration purposes such as wear resistant, corrosion resistant and porosity, lots of materials have been used in BPLC like aluminides of titanium, nickel and iron in form of X3Al (Levin, et al., 2003).

For experimental researches, numerical modelling will always be a good prediction for results since they are many factors affecting the cladding. Mazumder et. al (Mazumder, 2017) developed a self-consistent 3D transient model to examine the physical phenomena including heat transfer, melting/solidification phase change and liquid metal flow using a coaxial powder injection laser cladding process. The numerical results provided a good prediction to form a good metallurgical bonding by optimising the processing parameters. However, computational software and hardware still have limitations in simulating a model of accurate blown powder dynamics, melt pool morphology and melt pool fluid dynamic (Thompson, et al., 2015). Mechanical properties such as tensile, fatigue and impact testing are common applied to examine and evaluate the performance of the cladding structures.

Mechanical properties such as tensile, fatigue and impact testing are common applied to examine and evaluate the performance of the cladding structures. It’s found that DLD parts gave similar or higher values of properties compared to wrong and cast materials (Shamsael, et al., 2015).To better understand and investigate the BPLC process, high resolution microscopy of the generated metallurgies, micro-hardness test and tensile tests were completed as part of this work. Micro-hardness of laser cladding coupons varies significantly based on microstructural variations. Micro-hardness generally increases when microstructural variations decreases and keeps steady if majority of the structure is equiaxed when compared to the standard value for the material. Moreover, based on current novel low-power processing conditions, a majority equiaxed grain structure in stainless steel has been observed and thus potentially eliminating material anisotropy.2.

Moreover, based on current novel low-power processing conditions, a majority equiaxed grain structure in stainless steel has been observed and thus potentially eliminating material anisotropy.2. Experimental Work2.1 Experimental SetupA PRC 1.5kW CO2 laser was used with a TEM01 mode shape. The laser is focused and delivered via a central nozzle mounted on the z-axis which is perpendicular to the cladding X- Y traverse table. This CNC table is controlled by the computer programmes which can build designed structures through G-code. The moving speed and layer height can be altered inside the CNC G-code as well as shape the component. The powder is fed into the four nozzles (in157328th Annual International Solid Freeform Fabrication Symposiumred pipes, Figure 1) using a screw feed system on the top with Argon as assistant gas. The screw motor determines the powder mass.

The screw motor determines the powder mass. The gas flows for both powder feed and coaxial nozzles are controlled manually at a certain matching calibration to form an ‘ideal’ ‘Y’ shape of powder configuration (Figure 2 and Figure 3).Figure 2: Schematic of Four-port powder feed nozzle at Liverpool.Figure 3: 'Y' shape of powder configuration.Initial observations show that the process window is affected by build height, substrate choice, deposition material constraints or selection, powder delivery rates, laser power and process speed. Therefore, close control of process parameters is essential. The system has been specifically developed to give height control via this nozzle design and close control of very low powder delivery rate is possible with a custom design powder hopper.

The system has been specifically developed to give height control via this nozzle design and close control of very low powder delivery rate is possible with a custom design powder hopper. This nozzle can provide omnidirectional cladding with low angle powder feed and low velocity powder feed which helps to improve the catchment efficiency.Apart from the effects from power, processing velocity and powder materials, layer height of cladding builds is another essential factor which affects the microstructures and related mechanism. In terms of geometrical accuracy, numerous researchers have designed many control methods to detect layer height and specific measurements. The benefit from this nozzle design at Liverpool is an open loop system that can utilize the shape of the powder stream to control the layer height which can be simply altered from the G-code programs.

To investigate the effect of layer height, the height from 0.4mm to 1mm with 0.2mm segments were varied.The powder material used in this work is 316L stainless steel powder with the composition listed in Table 1.157428th Annual International Solid Freeform Fabrication SymposiumTable 1: Selected Austenitic Stainless Steel with chemical composition (nominal), wt. %.Alloy FeCCrNiMoSi MnS316LBalanced0.0316.0-10.0-2.0-1.02.00.0318.014.03.0Samples were sectioned, mounted and etched for analysis using optical microscope. To see the differences between single and multiple heating cycles of cladding builds, samples were selected from both intersected and non-intersected areas. Vickers hardness measurements were taken using a 500g load on the investigated surface. In terms of accuracy, five datum points were collected for each layer, equidistance throughout the sample.

Vickers hardness measurements were taken using a 500g load on the investigated surface. In terms of accuracy, five datum points were collected for each layer, equidistance throughout the sample. To explore the differences in mechanism between cladding and pure metal structures, a simple tensile test was taken.Samples were sectioned, mounted and etched for analysis using optical microscope. To see the differences between single and multiple heating cycles of cladding builds, samples were selected from both intersected and non-intersected areas. Vickers hardness measurements were taken using a 500g load on the investigated surface. In terms of accuracy, five datum points were collected for each layer, equidistance throughout the sample.

Vickers hardness measurements were taken using a 500g load on the investigated surface. In terms of accuracy, five datum points were collected for each layer, equidistance throughout the sample. To explore the differences in mechanism between cladding and pure metal structures, a simple tensile test was taken.2.2 Cladding Design and ProcessingTwo designed structures were laser clad and tested which both contained two intersecting circles of the same diameter. In terms of investigation on effect of multiple heating cycles and building methods, both intersected and non-intersected sections need to be taken into consideration when designing the geometries. Structure A (Figure 4 (a)) is intersected with a radius which built up with a linear method and laser indexing one layer by one layer.

Structure A (Figure 4 (a)) is intersected with a radius which built up with a linear method and laser indexing one layer by one layer. Structure B (Figure 5) two circles tangential which cladded in a helical manner (like infinity symbol ‘∞’) and the laser was constantly indexing upwards.(a1)(a2)Figure 4: (a1) Structure A with two circles intersected at half diameter; (a2) cladding path order.(b1)(b1)Figure 5: (b1) Structure B with two circles intersected at the edge; (b2) cladding path order.1575P0.04528th Annual International Solid Freeform Fabrication SymposiumBoth structures were laser clad under the same processing parameters which are shown in Table 2. All these parameters were kept constant during the cladding process.

All these parameters were kept constant during the cladding process. It is worth noting that the power used was only around 320W which was far less than the typical power of 800- 1000W for similar processes.Table 2: Consistent Processing Parameters of Structure A and B.Power (W)Process Velocity (mm/s)Powder Feed Rate(g/min)Coaxial Gas Flow(L/min)Powder Gas Flow (L/min)32042382.52.3 Cladding Design and Processing for Tensile TestAnother investigation into the mechanical property was achieved from tensile tests to obtain the ultimate tensile strength (UTS) and maximum load for the samples produced by laser cladding.

To generate the tensile test coupons, the orientation of the deposited tracks in relation to the direction of the tensile force was designed in two directions, one was perpendicular to the tensile force (structure C) and the other parallel (structure D) shown in Figure 6.(c)(d)Figure 6: (c) Structure C with vertical cladding and (d) Structure D with horizontal builds used for tensile test (Davis, 2006).Table 3 and Table 4 display the processing parameters and related dimensions for the two orientations. The parameters set were to explore the power effect and two types of powder catchment efficiency combined by power feed rate and gas flow rate.

The parameters set were to explore the power effect and two types of powder catchment efficiency combined by power feed rate and gas flow rate. All coupons were laser clad with 1mm layer height under cladding speed of 4 mm/s.Table 3: Processing parameters for structure C for vertical cladding tensile test samples and related dimensions(Figure 6 (c)).157628th Annual International Solid Freeform Fabrication SymposiumTable 4: Processing parameters for structure D with horizontal cladding tensile test samples and relateddimensions (Figure 6 (d)).3. Results and Discussion3.1 Microscopy Analysis 3.1.1 Structure A-- two circles intersected at half diameterIn terms of the non-intersected area, a typical directional solidification and epitaxial growth was observed in majority of cladding samples as well as strong microstructural variations.

Towards the outside of the sample the microstructure is fine cellular whereas towards the center the microstructure is a mixed columnar dendritic and relatively coarse cellular (Figure 7(b)). This will be due to the outside of the melt pool cooling at a faster rate than the center of the melt pool causing fine cells to form at the outer regions (Fu, et al., 2008). For the intersected part, due to the complex melting and re-melting process, in which multiple solidification fronts meet from cladding occurring in two separate directions, a complex microstructure is observed. This is comprised of several phases of columnar dendritic, epitaxial cells and fine equiaxed cells. From Figure 7 and Figure 8, the layer height of cladding is significant to the microstructures and its variations. Sample 4 (structure A) with 1 mm layer height provides most equiaxed structures but still suffers from different level of epitaxial growth from the solute bands.

Sample 4 (structure A) with 1 mm layer height provides most equiaxed structures but still suffers from different level of epitaxial growth from the solute bands. In addition to the variation based on the distance from the substrate, it is difficult to generate a general summary based on the observation from Figure 8.

In addition, the directional solidifications percentage was reduced by increasing the layer height, which can be seen from Figure 7 and Figure 8, due to the cooling rate of the cladding track that was affected.(a)(b)Figure 7: (a) Structure A with 0.4mm layer height intersected section, X100; (b) Structure A non-intersectedsection, X40.1577(a)28th Annual International Solid Freeform Fabrication Symposium(b)Figure 8: (a) Microstructure of Structure A non-intersection 8mm from substrate; (b) Microstructure ofStructure A with 1mm layer height non-intersection 18mm from substrate, x100 magnification.Figure 9: Overall view for the non-intersected region from the bottom to the top layers.Figure 10: Overall view for the intersected region from the bottom to the top layers.157828th Annual International Solid Freeform Fabrication Symposium3.1.2 Structure B -- two circles intersected at the edgeFigure 11 (a-c) contains the typical characteristic microstructures of the laser cladding process.

The solute banding from the solidification profile of the melt pool can be seen. The outside of the melt pool was cooling faster than the inside so that results in solidification directionality structures formed with epitaxial growth from the solute banding (Yue & Li, 2007). In other words, lots of dendritic structures propagating from solute bands in majority of layers of structure B. In the intersection part, majority of homogeneous cellular regions were concentrated with fewer variations in phases. Figure 11 (d) shows the equiaxed isotropic metallurgy observed under the low power conditions. This is a significant result for BPLC under this low power and related novel processing parameters.

Figure 11 (d) shows the equiaxed isotropic metallurgy observed under the low power conditions. This is a significant result for BPLC under this low power and related novel processing parameters. As equiaxed structures could eliminate material anisotropy.(a)(b)(c)(d)Figure 11: (a) (b) (c) Microstructure of Structure B with 1mm layer height intersection part on 6th, 9th, 12th layers from the substrate, with X40; (d) Microstructure of Structure B with 1mm layer height 11th layer intersected are zoom-in with X100 magnification.The results of the microscopy investigation indicated that three kinds of structures were generated: dendritic cells, equiaxed cellular and solidification directionality microstructures (Kurz, et al., 2001). The distributions of them were varied along the distance from the substrate.

The distributions of them were varied along the distance from the substrate. Through altering the process parameters, homogeneous and equiaxed microstructures were found with uniform distribution and in large amounts under these novel processing conditions. The essential control factors involved layer height, the distance from the substrate and the intersected points. These have significant effects on various microstructural distribution and growth tendency due to the multiple heating cycles.3.2 Micro-hardness Test 3.2.1 Structure A (Two circles intersected at half diameter)As can be seen from Figure 12, the micro-hardness of structure A varied from 170-240HV. The points were taken from the midpoints of the solute bands at various distances from the substrate.

The points were taken from the midpoints of the solute bands at various distances from the substrate. Compared to the typical hardness of wrought AISI 316L stainless steel ranging from157928th Annual International Solid Freeform Fabrication Symposium170-220HV (Kurgan, et al., 2012), the results achieved were very close which may indicate hardening does not occur during the cladding process. However, a higher hardness value was expected or hardening should exist as fine cellular microstructures were found; high thermal gradients would result in high thermal stress which would cause the hardness increase of the microstructures.

This discrepancy can be explained by the E/V values, as these were too low to cause the significant hardening effects in the material.Figure 12: HV hardness of Structure A with 0.4mm layer height (sample 1) and 0.8mm layer height (sample 3) both intersected and non-intersected area (320W, 4mm/s cladding speed, powder feed rate 3000rpm, coaxial and powder gas flow are 8 and 2.5L/min).Table 5: Micro-hardness and E/V value for Structure A both non-intersected and intersected area.SampleAverage Vickers’s Hardness (HV)E/V (J/mm3)1 (Non-intersected, 0.4mm/layer)189.5159.181I (Intersected, 0.4mm/layer)195.6159.182 (Non-intersected, 0.6mm/layer210.598.232I (Intersected, 0.6mm/layer)220.398.233 (Non-intersected, 0.8mm/layer207.671.313I (Intersected, 0.8mm/layer)223.471.314 (Non-intersected, 1mm/layer201.454.024I (Intersected, 1mm/layer)215.654.023.2.2 Structure B ()The micro-hardness for the structure was ranging from 190HV to 250HV (Figure 13), these were higher than both structure A and the AISI standard.

This indicates that the material158028th Annual International Solid Freeform Fabrication Symposiumhardening occurred. This can be explained as a lower variation of microstructures and a smaller distribution of directionally solidification structures was observed.

In addition, this indicated that the existence of equiaxed cellular structures would help to increase the hardness properties which concluded by Wu et al (Wu, et al., 2010).Figure 13: HV hardness of Structure B with 0.4mm layer height (sample 5) and 1mm layer height (sample 8) both intersected and non-intersected area (320W, 4mm/s cladding speed, powder feed rate 3000rpm, coaxial and powder gas flow are 8 and 2.5L/min).Table 6: Micro-hardness and E/V value for Structure B both non-intersected and intersected area.SampleAverage Vickers’s Hardness (HV)E/V (J/mm3)5 (Non-intersected, 0.4mm/layer)205.5162.075I (Intersected, 0.4mm/layer)210.6162.076 (Non-intersected, 0.6mm/layer)210.5101.586I (Intersected, 0.6mm/layer)225.6101.587 (Non-intersected, 0.8mm/layer)211.373.677I (Intersected, 0.8mm/layer)234.173.678 (Non-intersected, 1mm/layer)242.156.158I (Intersected, 1mm/layer)254.256.153.3 Tensile TestThe structures C and D (shown in Figure 6) were subjected to a tensile test.

From the results generated in Table 7 and Table 8, it can be summarized that the horizontal samples which the track orientation was parallel to the tensile force show higher maximum breaking load and ultimate tensile stress than perpendicular ones.158128th Annual International Solid Freeform Fabrication SymposiumTable 7: Averaged mechanical properties for vertical samples.SampleMax.Load (kN)Ultimate Tensile Stress (MPa)Elongation (%)C1 (320W)12.15294.2613.55C2 (400W)0.7417.566.96C3 (600W)7.00156.636.55C4 (800W)9.10175.287.97Table 8: Averaged mechanical properties for horizontal samples.SampleMax.Load (kN)Ultimate Tensile Stress (MPa)Elongation (%)D1 (320W)15.71389.7338.88D2 (400W)14.05365.1045.00D3 (600W)13.04274.3830.37D4 (800W)15.38335.4629.75The cladding tracks perpendicular to the tensile force show a lower strength as the interfaces were weaker compared to the solid track area which means there are easier to be pulled apart from another track with weaker bonding (Dahotre, 1998).

As discussed, directional solidification grains will grow from the solute band in different proportions based on changing processing parameters. In terms of microstructures under this condition, the directional solidification grains grew the same direction as the applied tensile force which allowed fewer barriers to dislocation movement; similar with being deformed or breaking easier and therefore a low mechanical strength. The failure mostly occurred from the track boundaries (Figure 14158228th Annual International Solid Freeform Fabrication Symposium(a)). For the sample with track direction along with the tensile force, higher strength and stiffness were achieved. The solute boundaries retained mechanical integrity and coherence during the pulling process. Moreover, the mobility of grains dislocations can determine the different degrees and direction of deformation.

Moreover, the mobility of grains dislocations can determine the different degrees and direction of deformation. If a majority of equiaxed structures could be realized this would provide more barriers to prevent distortion and improve material anisotropy. From the final failure condition shown in Figure 14 (b), it can be seen that the failure path agreed with the solidification grains from different layers.Figure 14: (a) Vertical build which tracks perpendicular to the tensile force; (b) horizontal build which tracks parallel to the force.4. ConclusionsBased on the current work, it can be concluded that novel processing parameters have been found that provide better distributions of equiaxed structures. Due to the uniform distribution of homogeneous equiaxed structures, the structure B tested stronger in mechanical performance.

Due to the uniform distribution of homogeneous equiaxed structures, the structure B tested stronger in mechanical performance. In terms of tensile testing, structure D displayed a higher breaking load which related with cladding bonding efficiency and uniformity of microstructures. Processing powers of approximately 320W were used to generate these structures, this is significantly less than previous work in this area.A limitation for laser-assisted blown powder deposition process in many applications is that the material properties of parts made tend to show anisotropy due to directional solidification of the microstructures as the re-melting conditions of the multi-layer process. This is not a problem for some applications such as an aero engine fan blade, but such anisotropy leads to severe design constraints in other AM components resulting in a requirement for post-process heat treatment.

BPLC exploration in this work has identified novel processing conditions that yield equiaxed grain structures in stainless steel 316L and thus potentially eliminate material anisotropy. This is potentially a step change in the capability of AM.Future work will confirm these findings in different materials and the variations of microstructures according to the distance from the substrate still need to be further explored. The process window is affected by build height, substrate choice, deposition material constraints or selection, powder delivery rates, laser power and process speed. Close control and optimisation of all these parameters will also form the basis for future work.158328th Annual International Solid Freeform Fabrication SymposiumReferences [1] Dahotre, N. B., 1998. Lasers for Metallic and Intermetallic Coatings. In: Lasers in Surface Engineering. Tullahoma: ASM International ,p. 141.[2] Davis, S. J., 2006.

ND CHARACTERIZATION IN VERY HIGH POWER ULTRASONIC ADDITIVE MANUFACTURING M.R. Sriraman*, Hiromichi Fujii*, Matt Gonser*, S.S. Babu*, and Matt Short+ *Department of Materials Science and Engineering, The Ohio State University, Columbus, OH 43221, USA +Edison Welding Institute, Columbus, OH 43221, USAAbstract Solid parts were produced by Very High Power Ultrasonic Additive Manufacturing (VHP-UAM) at room temperature using 150 μm thick tapes of 6061 aluminum and 110 copper alloys. Processing was done at 20 kHz frequency over a range of parameters (26 –36 μm vibration amplitude, 5.6 – 6.7 kN normal force, and 30.5 - 35.5 mm/s travel speed). Softening of materials (up to about 14% in 6061 Al and 23% in 110 Cu) was noted facilitating enhanced plastic flow and a reduction in interfacial voids.

Softening of materials (up to about 14% in 6061 Al and 23% in 110 Cu) was noted facilitating enhanced plastic flow and a reduction in interfacial voids. Evolution of fine recrystallized grains (0.3-4 μm in 6061 Al and 0.3-10 μm in 110 Cu) from an initial coarser grain structure (up to 8 μm in 6061 Al and 25 μm in Cu) was observed at the build interface regions. Bonding between layers in both materials seems to have occurred by dynamic recrystallization and movement of grain boundaries across the interface. The energy required for the above physical processes is derived from interfacial adiabatic plastic deformation heating.IntroductionAdditive manufacturing is a set of manufacturing processes that involves the productionof a 3-D part layer by layer directly from a CAD model [1]. Also known as layered manufacturing or rapid prototyping, this technology enables building of complex shaped parts that have hitherto been impossible by traditional means [1].

Also known as layered manufacturing or rapid prototyping, this technology enables building of complex shaped parts that have hitherto been impossible by traditional means [1]. Ultrasonic Additive Manufacturing (UAM) is one such process used to fabricate parts of near net shape from metal tapes [2]. Based on ultrasonic seam welding, this solid state joining process involves successive bonding of metal tapes one over the other [3]. Ultrasonic vibrations (20 kHz frequency) are applied laterally to the foil through a sonotrode under a static normal force as the sonotrode rolls along the length of the tape. The scrubbing that takes place between the tapes ruptures the oxide layers between them, promoting nascent metal-metal contact and eventually their bonding. The sonotrode surface is usually textured to enable gripping of the foil and facilitate welding of the faying surfaces. This leaves a rough imprint on the top-surface of the tape.

The sonotrode surface is usually textured to enable gripping of the foil and facilitate welding of the faying surfaces. This leaves a rough imprint on the top-surface of the tape. These imprints affect the bonding of the subsequent layer. The main process parameters for a given sonotrode texture are vibration amplitude, static normal force, sonotrode travel speed and pre-heat temperature [3].A number of investigations have been carried out to characterize the bonding betweenmetals in UAM [4-8]. Microstructural changes occurring under this process and the ensuing mechanical properties (hardness and strength) have been studied [4-11]. However, due to limited power capability of previous systems [12], adequate bonding has not been possible even in Al despite the pre-heat [3, 13]. This results in a large number of interfacial voids and discontinuities throughout the build [3, 13]. Furthermore, welding of materials harder than aluminum has not been possible.

This results in a large number of interfacial voids and discontinuities throughout the build [3, 13]. Furthermore, welding of materials harder than aluminum has not been possible. Preliminary attempts have been made to bond copper and some dissimilar metal combinations [6-7] with limited success. These shortcomings in the current UAM method have led to the development of very high power ultrasonic additive manufacturing (VHP-UAM). Such a process will involve higher ultrasonic power (or energy) realized through larger vibrationamplitudes and normal force levels. Such a VHP-UAM machine has been designed and developed by the EWI [12]. It is capable of delivering up to 9 kW of power at 20 kHz resonant frequency through the use of two transducers operating in tandem on either side of the sonotrode [12, 14]. A schematic of this is shown in Fig. 1. The two transducers are set to vibrate 180° out of phase, in “push-pull” mode.

A schematic of this is shown in Fig. 1. The two transducers are set to vibrate 180° out of phase, in “push-pull” mode. The transducer displacements are reinforced resulting in vibration amplitudes of up to 52 µm (from 26 µm in the conventional machine) [12]. In addition, the machine also has the capability for higher static normal forces – up to 15 kN [much higher than in the conventional machine (2 kN)] [12]. In this paper, 6061 Al (Al- 1% Mg, 0.6% Si, and 0.3% Cu by weight) [15] and 110 Electrolytic Tough Pitch Cu (99.95 % Cu and 0.04% O by weight) [15] have been processed by VHP-UAM. The bonds were then characterized by a variety of techniques.Fig. 1 Schematic of the double transducer-sonotrode system used in the VHP-UAM machine for enhanced vibration amplitudes.Experimental Procedure Separate builds of 6061 Al and 110 Cu were made by welding of tapes 150-µm thick successively one over the other. The first tape was always laid on (welded) to a 12-mm thick 3003Al-H14 base plate.

The first tape was always laid on (welded) to a 12-mm thick 3003Al-H14 base plate. The approximate dimensions of the builds were 25.4 mm width X 170 mm length X 2 mm height. Processing of materials was done at room temperature using the parameters shown in Table 1.Table 1 VHP-UAM processing parameters used in the current investigationsMaterial 6061 Al-H18 110 Cu-HardVibration amplitude (μm) 26 36Normal force (kN) 5.6 6.7Travel speed (mm/s) 35.5 30.5The builds were characterized on transverse sections perpendicular to the weldingdirection. The samples were sectioned at low speed using a diamond blade and prepared in accordance with standard metallographic practice. They were then final polished with 0.02 µm colloidal silica on a vibratory polisher for ~36 hours. Optical microscopy and Scanning electron microscopy (SEM) (using the Quanta 200 Scanning electron microscope) were used to observe un-bonded regions and interfacial discontinuities.

Optical microscopy and Scanning electron microscopy (SEM) (using the Quanta 200 Scanning electron microscope) were used to observe un-bonded regions and interfacial discontinuities. Microhardness mapping across the build (grid comprising of 31 rows spaced at 60 μm intervals and 20 columns spread over a length of about 20 mm) was done using a Leco AMH-43 Microhardness analyzer. The microstructure evolutionat the interface regions was investigated employing the technique of Orientation imaging microscopy (OIM) on Quanta 200 under the following conditions: 20-25 kV, spot size 6, 21-25 mm working distance, and 0.1-0.2 μm step size. For comparative studies, the original base material foils were also characterized on transverse sections. Results and Discussion6061 Aluminum Alloy BuildsTypical optical images from an as-polished 6061 Al build sample is shown in Fig. 2. Thepresence of a few voids is seen in the sample center (Fig.

Results and Discussion6061 Aluminum Alloy BuildsTypical optical images from an as-polished 6061 Al build sample is shown in Fig. 2. Thepresence of a few voids is seen in the sample center (Fig. 2a) while relatively more voids are observed towards the edges (Fig. 2b). The void density is estimated to be 0.2 - 0.4% in a total area of 2 mm2. These un-bonded regions however are noticeably lower than what is typically seen in conventional UAM (Fig. 2c). Although the processing parameters used here are not optimal, the observation of a low void density indicates that improved bond quality is possible from the VHP process.Fig. 2 Optical images of as-polished 6061 Al build sample (transverse section) showing (a) the void distribution in center and (b) edge. The optical image of the sample made by conventional UAM is shown in (c) for comparison [13].

The optical image of the sample made by conventional UAM is shown in (c) for comparison [13]. A schematic of the transverse section showing regions corresponding to center and edge (denoted by C and E respectively) is also shown for reference.The presence of more voids towards the edges could be attributed to a non-uniform contact pressure distribution at the interfacial region when a layer is bonded. Imprints of the sonotrode surface texture left on the foil (Fig. 3) show regions close to the edge of the tape (Fig. 3a) that are left “untouched” by the sonotrode surface. The tape center (Fig. 3b) on the contrary has a more uniform surface texture (Fig. 3b).Fig. 3 Optical images of 6061 Al build top surface showing regions in the edge (a) where the sonotrode does not make contact unlike in the center (b).Fig. 4 Hardness map over the 6061 Al build sample showing a decrease in hardness in comparison to that of the original foil.Microhardness measurements (Fig.

4 Hardness map over the 6061 Al build sample showing a decrease in hardness in comparison to that of the original foil.Microhardness measurements (Fig. 4) revealed softening of the material, at both theinterface and bulk regions, in comparison to the hardness of original foil (78 HV). A hardness drop of up to ~14% is noted. However, as can be seen (Fig. 4), the extent of softening is not uniform across the build, with certain regions being almost as hard as the original foil. Such “hard” regions are found to be distributed throughout the build (Fig. 4). Although the reason for this is not clear, it is hypothesized that they could correspond to regions where plastic flow has not been homogeneous. Such inhomogeneous flow can arise from a non-uniform contact pressure distribution beneath the sonotrode [16]. Also, no specific trend is noticed with respect to the hardness variation either across the build height or along the transverse direction (Fig.

Also, no specific trend is noticed with respect to the hardness variation either across the build height or along the transverse direction (Fig. 4).Microstructure analysis shows the evolution of a fine equiaxed grain structure at the interface region (Fig. 5). Fig. 5b is an inverse pole figure (IPF) from one such region in the build sample. Regions of the same color signify those grains that have their corresponding crystallographic planes (indicated by the color triangle) oriented parallel to the sample surface.The IPF of the original foil is shown in Fig. 5a. Although the confidence index has been rather low (0.14), the difference between the microstructures is clear. It is seen that a larger pancake shaped grain size distribution in the original foil (up to 8 µm) has been reduced to a finer equiaxed grain structure (0.3-4 µm) at the interface region following bonding (Fig. 5c).

5c). The finer grain structure that is evolved during processing also has a larger fraction of grain boundaries with high angle misorientations (15-65°) between them (Fig. 5d). The formation of such fine equiaxed grains with large misorientations suggests the occurrence of recrystallization. Since the interface is not distinctly visible (Fig. 5a), the grain boundaries (of recrystallized grains) must have moved across the interface, leading to metallurgical bonding.Fig. 5 Inverse pole figures of (a) the original 6061 Al foil and (b) an interface region from the build. Analyses of the data shows (c) finer grain size distribution and (d) larger grain misorientations at the build interface.A clear change in the crystallographic texture is also noted as seen from the (111) polefigures (Fig. 6). The original foil (Fig. 6a) shows a typical {112} <111> “copper” texture that is commonly observed with heavily cold rolled face centered cubic metals [17].

6). The original foil (Fig. 6a) shows a typical {112} <111> “copper” texture that is commonly observed with heavily cold rolled face centered cubic metals [17]. The interface region however seems to display a weak “cube” texture ({001}<100>) [17] again pointing to recrystallization of the initial deformed structure. This is shown in Fig. 6b. Recrystallization has been observed earlier in conventional UAM of Al (3003 alloy) [3, 18]. In the current investigations however, the phenomenon observed can only be attributed to dynamic recrystallization (DRX) since no external heating is involved. DRX signifies recrystallization of grains even as it is being deformed. Such a phenomenon has also been reported in conventionalUAM of 6061-O Al done without pre-heat [10]. The occurrence of DRX probably explains the softening seen in our investigations.Fig.

Such a phenomenon has also been reported in conventionalUAM of 6061-O Al done without pre-heat [10]. The occurrence of DRX probably explains the softening seen in our investigations.Fig. 6 (111) Pole figures of (a) the original 6061 Al foil and (b) an interface region of the build sample showing a weak cube texture evolution in the interface from an initial copper texture.(a)110 Cu Alloy BuildsBonding in 110 Cu is seen to comprise of the formation of wavy interfaces with evidenceof turbulent plastic flow in certain regions (Fig. 7). The presence of such interfacial instabilities was clearly brought out in the etched build sample under the SEM (Fig. 7). This could be an effect of the material flow occurring along the ridges and valleys between the contacting layers during processing.

7). This could be an effect of the material flow occurring along the ridges and valleys between the contacting layers during processing. The wavy nature of interfaces is probably connected with the ability of copper to work soften and flow easily (than Al) with increasing temperatures, due to its lower stacking fault energy (SFE) [19-21]. It is also apparent (Fig. 7) that the tapes had individually thinned down by about 5% from the original thickness of 150 µm.Fig. 7 Secondary electron images of 110 Cu build sample etched with 50% nitric acid.Softening of the material throughout the build (even greater than in 6061 Al) is observed here again. Fig. 8 shows the hardness map over the build sample. In comparison with a harder original foil material (108 HV), the build layers are seen to have softened up to ~23% (Fig. 8). However, there are some “hard” regions as well. Again, no specific trend is seen in the hardness variation along the build height or in the transverse direction (Fig.

8). However, there are some “hard” regions as well. Again, no specific trend is seen in the hardness variation along the build height or in the transverse direction (Fig. 8).Fig. 8 Hardness map over the 110 Cu build sample showing a decrease in hardness in comparison with the original foil.The evolution of a fine equiaxed grain structure at the interface region was again noted (Fig. 9). Figures 9a and 9b show the IPF of the original foil and an interface region in the build sample. The difference between the two images (Figs. 9a and b) is striking with the latter showing evidence of highly localized shear (Fig. 9b). Grain size analysis reveals that a relatively coarse grain structure of up to 25 µm size in the as-received foil is reduced on processing to a more uniformly finer grain size ranging between 0.3-10 µm (Fig. 9c).

9c). This fine grain structure in the interface region also contains a larger fraction of grains with high angle misorientations (15- 60°) between them than the original foil (Fig. 9d) pointing once again to recrystallization and consequent metallurgical bonding. Again, with no external heating involved in the processing of the build, the evolved grain structure at the interface region could be attributed only to DRX. It is probably the first time this phenomenon is being observed in UAM of Cu. A change in texture is seen here too - between the original foil and the interface region of the build – as seen in the (111) pole figures (Fig. 10). It is interesting to note that a deformation texture (copper type) seems to evolve following processing (Fig. 10b). This could be associated with the shearing seen therein (Fig.

10). It is interesting to note that a deformation texture (copper type) seems to evolve following processing (Fig. 10b). This could be associated with the shearing seen therein (Fig. 9b) and is being further investigated.With the observation of dynamic recrystallization occurring in both materials at the interface, it appears the starting cold worked coarse grain structure of the original foils has gone through the stages of further plastic deformation (expected to be more in Cu due to its high work hardening rate) and dynamic recovery (in Al) during VHP-UAM. Al alloys, with their high SFE, are known to exhibit continuous dynamic recrystallization (CDRX) [10, 22] especially under hot working/ thermomechanical processing [19-21] where the subgrains undergo a continuous rotation until they form high angle boundaries between them [10, 22].

Copper, on the other hand, being a material with medium SFE, is known to exhibit discontinuous dynamic recrystallization (DDRX) under hot working conditions [23]. This involves the nucleation of grains with high angle misorientation with the parent grains [23].Fig. 9 Inverse pole figures from (a) the original 110 Cu foil and (b) an interface region in the build sample. Analyses of the data shows (c) finer grain size distribution and (d) larger grain misorientations at the build interface.Fig. 10 (111) Pole figures of (a) the original 110 Cu foil and (b) an interface region of the build sample showing the possible evolution of a copper texture at the interface region.(b)The occurrence of recrystallization suggests that the interfacial temperature during the process must have reached close to the homologous temperature of 0.5 (although in Al alloys, the temperature could be lower) [19].

It is further interesting to note that this might have actuallyoccurred within a few milliseconds only [24] entirely due to interfacial heating. The evolution of recrystallized grains at the interface explains the softening observed in both materials. The heat generated at the interface could also get dissipated into the bulk. Since the starting material is in the cold rolled condition, such heat dissipation could lead to recovery-like effects in the bulk of the layers [25]. This probably explains the softening through the thickness of the layers.Since the process involves shearing of small micron sized asperities between the foils at afrequency of 20 kHz and high vibration amplitudes, very high strains and strain rates can develop. Strain rate estimates under ultrasonic welding/ UAM point to a figure as high as 104 – 105 s-1 [7, 26]. This could result in high strain rate deformation heating under adiabatic conditions resulting in large temperature excursions [27-28].

This could result in high strain rate deformation heating under adiabatic conditions resulting in large temperature excursions [27-28]. The flow stress of the material is thus decreased appreciably even as the high strain rate tries to increase it [20]. Between these competing processes, the effect of temperature eventually prevails resulting in work softening through dynamic recovery (DRV) and/or recrystallization (DRX) [17, 20-21]. Such a behavior can induce plastic instability in the material promoting enhanced flow, more so in materials with relatively lower SFE [17, 21]. This is probably the reason for the turbulent flow in Cu. Such inducement of enhanced plastic flow can result in a more effective collapse of asperities too. This would minimize voids and un-bonded regions, as was noted in this research. A further reduction in void density (than reported here) was also seen to be possible under VHP processing (as part of our overall research program) [25].

A further reduction in void density (than reported here) was also seen to be possible under VHP processing (as part of our overall research program) [25]. Temperature rise estimates made based on adiabatic heating during metal working show that the excursions in temperature over every cycle can indeed be large and close to a homologous temperature of 0.5 [24]. VHP-UAM can thus be compared to hot working (albeit in a much localized fashion) where such temperatures are involved promoting DRX [19].Based on the above, the following bonding mechanism is hypothesized in VHP-UAM:1. Scrubbing of layers leading to oxide film rupture between all contacting asperities, their delamination, and dispersion - due to differential deformation characteristics between the oxide and the metal.2. Contact of nascent metal asperities, their interfacial cyclic shearing and strain hardening. 3.

Contact of nascent metal asperities, their interfacial cyclic shearing and strain hardening. 3. Formation of hysteresis loops, plastic deformation heating under adiabatic conditions due to high strain rates.4. Rapid localized temperature rise with every cycle together with plastic strain accumulation.5. Softening by DRV/ DRX and decrease in flow stress and modulus of the interfacial metal.6. Enhanced plastic flow, asperity “collapse”, and further softening. 7. Mutual migration of boundaries of recrystallized grains across the interface. 8. Formation of a series of “micro welds”.Although listed sequentially, these stages in bonding could occur concurrently. Forinstance, the initial plastic deformation cycles could lead to adiabatic heating conditions, which in turn facilitate more plastic flow. Again, the softening that takes place by the mechanism of DRV/ DRX could promote more plastic flow and a possible further softening due to continued heating.

Again, the softening that takes place by the mechanism of DRV/ DRX could promote more plastic flow and a possible further softening due to continued heating. Also, even as the contacting asperities collapse against each other, new asperities start coming into contact as the cycle repeats itself.In order to test the above hypothesis of plastic deformation heating being the driver toDRV/ DRX in VHP-UAM, some temperature measurements using K type thermocouples were made in Al alloys (3003, 6061) and 110 Cu [29]. A higher temperature rise was recorded in the higher strength materials namely 6061 Al and 110 Cu in comparison with 3003 Al under similar processing conditions [29] demonstrating the effect of plastic deformation heating. The observations also fit into the phenomenological model of adiabatic temperature rise due to plastic work [24]. Based on the above mechanism, bonding in a given material must also depend upon the energy that is put in through the processing parameters.

Based on the above mechanism, bonding in a given material must also depend upon the energy that is put in through the processing parameters. This input energy will in turn determine the extent of temperature rise at the interface promoting bonding through the mechanism discussed. This phenomenon was again verified by temperature measurements in 3003 Al [29]. Although all the three parameters - vibration amplitude, normal force, and travel speed were found to influence the temperature rise [29], the effect of amplitude was the most pronounced [29]. These results also agree with the phenomenological model [24]. According to the model, the temperature rise will be a function of the shear strain associated with plastic work. Since the shear strain is proportional to the amplitude [24], higher the amplitude, larger will be the temperature increase. The results [29] of increased temperature rise with amplitude (in 3003 Al) corroborates with observations made earlier [14].

The results [29] of increased temperature rise with amplitude (in 3003 Al) corroborates with observations made earlier [14]. A higher degree of bonding, as assessed by qualitative peel testing, was noted under increased amplitudes [14].Future work would involve mechanical testing of VHP-UAM builds to evaluate the“bond strengths” and identify optimal processing parameters for a given material. This would confirm the above hypothesis of temperature rise being a measure to the bond quality. Also, it would be interesting and important to observe the microstructural evolution away from the interfaces to rationalize the softening in the bulk.Conclusions Multilayer builds of 6061 Al and 110 Cu were made by very high power ultrasonicadditive manufacturing at room temperature.

With enhanced plastic flow facilitated through more intense processing parameters than in conventional ultrasonic additive manufacturing, voids/ unbonded regions are significantly low (only about 0.2-0.4% in 6061 Al over 2 mm2). Wavy interfaces with features of turbulent plastic flow are noticed in 110 Cu. Bonding in both materials is seen to take place through the process of dynamic recrystallization and a possible migration of grain boundaries across the interface. While the grain structure in 6061 Al is reduced to 0.3-4 μm from an initial starting foil grain size of up to 8 μm, it is refined to 0.3-10 μm from a grain size of up to 25 μm in the starting foil. This is accompanied by a change in crystallographic texture too at the interface region. Softening of materials from the original foil hardness is also observed (up to about 14% in 6061 Al and 23% in 110 Cu). This is probably responsible for the enhanced flow.

Softening of materials from the original foil hardness is also observed (up to about 14% in 6061 Al and 23% in 110 Cu). This is probably responsible for the enhanced flow. The differences in the extent of softening between 6061 Al and 110 Cu could be attributed to the difference in their stacking fault energies. The temperature rise promoting dynamic recrystallization during bonding seems to occur through plastic deformation heating between contacting asperities. The above results illustrate the potential benefit of VHP-UAM.Acknowledgements The authors are thankful to the Ohio department of development for funding this research work through the Third frontier Wright Projects program. The guidance and support received from Dr.Karl Graff of EWI throughout this research is greatly appreciated. Kind assistance from the staff/ students of EWI/ OSU is also gratefully acknowledged.References[1] Xiaochun Li, Hongseok Choi, Yong Yang, Thin Solid Films 420-421 (2002) 515. [2] Dawn R.

nding Methods for Laminated ToolingbyDaniel, F. Walczyk, Nathanael Y. Dolar Department of Mechanical Engineering, Aeronautical Engineering and Mechanics Rensselaer Polytechnic Institute Troy, New YorkAbstractLaminated tooling consists of an array of stacked laminations that are mechanically clamped or bonded together, depending on the requirements of the manufacturing process. Various manufacturing processes that can benefit from tooling constructed of laminations include sheet metal forming, thermoforming, composites molding, metal extrusion, injection molding, resin transfer molding, and compression molding. When bonding of the laminations is required (e.g., incorporation of conformal cooling passages for injection molding temperature control) then laminations can be joined together by diffusion bonding, brazing and using adhesives. However, for a tooling engineer to effectively design a laminated tool, the physical and mechanical properties of these joints must be known.

However, for a tooling engineer to effectively design a laminated tool, the physical and mechanical properties of these joints must be known. Consequently, a set of experiments is outlined for determining the tensile, shear, and peel strengths, tensile and shear elastic moduli, thermal contact resistance, and specific permeability (for gasses or liquids) of the aforementioned bonded joints for both steel and aluminum laminations. Some preliminary results with aluminum and future work are presented.IntroductionThe concept of making intricate parts (e.g. transformer core, door lock assembly, gear) by assembling profiled or contoured laminations is a well-established manufacturing. technique [15]. Laminated constructions have also been investigated, and successfully implemented by a few u.S. and Japanese companies [1,2] as a direct rapid tooling technique because of demonstrable advantages over other more conventional methods like CNC-machining and EDM.

and Japanese companies [1,2] as a direct rapid tooling technique because of demonstrable advantages over other more conventional methods like CNC-machining and EDM. These advantages include the elimination of tooling accessibility problems, reduced limitations on die geometry, faster fabrication, and easier incorporation of conformal cooling passages [1,3]. However, laminated tooling techniques have not yet seen widespread use in industry. One of the main reasons for industry's hesitance to adopt laminated tooling is the problem of having to make a rigid tool out of an array of laminations. As shown in Figure 1, the two methods for accomplishing this are by mechanical means (i.e., clamping laminations together) and by bonding methods (i.e., welding, brazing, soldering, diffusion bonding, and adhesives).This paper focuses on the latter method (i.e., bonding), especially for laminations with beveled edges as shown in Figure 1.

Furthermore, it will address the problem of bonding laminated tooling in a more comprehensive manner than has been done previously. Consequently, previous work in this area will first be discussed. A survey of manufacturing processes that can benefit from using laminated tooling will also be mentioned along with process. requirements that are needed for' tooling design. This will be followed by discussion of common bonding techniques available for laminations made out of the two most common tooling211materials used by industry, i.e., aluminum and steel. The complete experimental procedure of an ongoing research effort to characterize the mechanical and physical properties (i.e., tensile, shear and peel strengths, tensile and shear elastic moduli, thermal contact resistance, permeability, and dimensional changes) of bonded steel and aluminum laminations will be outlined.

When this is the case, then the laminations will need to be bonded together into a solid die in order to meet these requirements (e.g., sealing required for conformal cooling passages, Some preliminary work in lamination bonding includes laser welding lamination edges [7,8], brazing laminations together [9], bonding paper laminations together with a heated polymer layer [10], using adhesives for bonding plastic laminations [11], cementing steel laminations with a combination of adhesives and edge welding [12], diffusion and pressure bonding steel laminations together [13], and tack welding adjacent layers with an Nd:YAG laser [14]. Unfortunately, none of this previous work has investigated the specific mechanical and physical requirements (i.e., tensile, shear & peel strengths, tensile and shear elastic moduli, thermal contact resistance and permeability at various temperatures) for bonded joints that's required for a particular manufacturing process.

Nor have any automated methods for bonding tool laminations (except for Ref. 10) been developed.increased shear strength for sheet metalforming).212Candidate Manufacturing Processes for Bonded Laminated ToolsBy virtue of previous research and a need for enhanced die or mold performance over that the following manufacturingfabrication means, which is achievable using conventional processes have been chosen as candidates for laminated tooling:Sheet Metal Fonning including cold and hot matched-die fonning, drawing, and hydrofonning. fonning of sheet metal over a fonn die(s) of the desired shape. The metal is stretched to approximately 2% to 4% total strain, just beyond its yield point, to retain the contour of the die.

fonning of sheet metal over a fonn die(s) of the desired shape. The metal is stretched to approximately 2% to 4% total strain, just beyond its yield point, to retain the contour of the die. High nonnal and shear forces on the die surface will tend to peel and shear bonds between laminations.Thennofonning involves heating of a polymer sheet to its softening point, above the material's glass transition temperature (55-90°C), and drawing the pliable material against the contours (using vacuum or applied pressure) of the mold where it is held until cooled. Mild nonnal forces will tend to shear and pull bonds between laminations.Composites molding involves curing of composite materials using a male or female mold at room temperature (25°C) and pressure or higher (autoclave curing). The pressures imposed by the autoclave is an overall external pressure, in essence, hydrostatic. Although in some cases, there is shrinkage involved with the heating of the composite.

The pressures imposed by the autoclave is an overall external pressure, in essence, hydrostatic. Although in some cases, there is shrinkage involved with the heating of the composite. This can lead to some shearing loads on lamination bonds. Confonnal cooling passages are sometimes used in composites molding.Metal Extrusion involves forcing a billet of material (Aluminum, copper, magnesium, zinc, tin and their alloys) through a die opening, by way of a ram or a press, to produce a desired cross-sectional shape. Hot extrusion requires that the tool must be able to operate under severe conditions of temperature, pressure, and abrasive wear. Extrusion of metals creates very high shearing and nonnal forces on the inside surface of the billet container and die. Consequently, the lamination bonds will be subjected to tensile, shear and peel loads, regardless of lamination orientation.

Consequently, the lamination bonds will be subjected to tensile, shear and peel loads, regardless of lamination orientation. Cooling of the extrusion die with confonnal cooling passages is not a common practice because of the difficulty of making these passages using conventional machining practices.Injection molding involves melting and pressurizing granular or powdered thennoplastic polymer in a separate chamber and then forcing the molten polymer into a relatively cooler mold by way of a runner and injection system. It is in this "cooler" mold cavity that the part solidifies and fonns. The high nonnal forces on the inside of the mold cavity (resulting from the high injection pressures) create very high tensile and peeling loads on the lamination bonds.

The high nonnal forces on the inside of the mold cavity (resulting from the high injection pressures) create very high tensile and peeling loads on the lamination bonds. Confonnal cooling passages are commonly used in injection molding dies.Resin transfer molding involves infusing liquid resin (i.e., thennosetting polymer) by vacuum, applied pressure or both into a pre-fabricated/assembled dry fiber prefonn contained by a mold. This molding process heats the granulated or powdered thennosetting polymers and while in the softened state, forced into the mold cavity through a runner system. The mild nonnal forces on the inside of the mold cavity create tensile and peeling loads on the lamination bonds.

The mild nonnal forces on the inside of the mold cavity create tensile and peeling loads on the lamination bonds. The mold is preheated to help in fonning and curing thereby requiring confonnal cooling passages.Compression molding involves compressing a precise amount of thennosetting polymer or elastomer between heated mold halves and forcing the material to flow and confonn to the shape of the mold cavity. Heating of the dies causes the thennoset to polymerize and cure into a solidified part. The nonnal forces on the inside of the mold cavity create tensile and peeling loads on the lamination bonds. Confonnal cooling passages can be used to preheat the mold halves.Important process information on these candidate processes for consideration with internal or external pressuresbonded joints include the maximum operating temperatures,213involved, thermal conductivity requirements, permeability requirements, and typical die materials used.

lowlowhighhighhighhighmedium to highPermeability Requirementsnonevacuum, conformal cooling coolingselective heating & cooling vacuum, heatingcoolingvacuum, heatingApplicable Bonding Methodsdiffusion, brazing or adhesives diffusion, brazing, or adhesives 00 diffusion, brazing diffusion, brazing, or adhesives 00 diffusion, brazing, or adhesives 00 diffusion, brazing, or adhesives 00 diffusion, brazing or adhesives** *** + ++ o o 00uses both thermosettmg (TIS) and thermoplastiC (TIP) polymers but generally TIP uses TIP only uses TIS only uses both TIP and TIS with equal efficiency as in fiberglass based composite polymers along with various graphite based epoxies and TIS depending on the material under extreme conditions, though usually hydrostatic at 2.4 MPa adhesives to some extent depending on the pressures necessary to form the material properlyLamination Bonding Methodsto steel, and aluminum to aluminum including diffusion bonding, brazing, and joining with adhesives [16].

Each of the aforementioned bonding methods is discussed below:There are several methods suggested for bonding steelDiffusion bonding involves the filling of voids (asperities) at a bond interface by migration of molecules. Specifically, the diffusion kinetics at the bond are accelerated by applying the appropriate pressure and temperature to the lamination joint over a period of time. For example, Nakagawa et al. [13] diffusion bonded 55 sheets of cold-rolled steel by heating them together in a vacuum at a temperature l100°C followed by the application of pressure of 5.9 MPa for one hour. Advantages of this process include no visible bond line, a clean fluxless process, strength of the bond up to 100% of the parent material strength, and no significant thermal contact resistance.

Advantages of this process include no visible bond line, a clean fluxless process, strength of the bond up to 100% of the parent material strength, and no significant thermal contact resistance. Disadvantages include the limited application of this process to specific combinations of materials (e.g., aluminum to aluminum doesn't work), mild shrinkage (400 microinches at a joint), need for a hot press, need for a material that forms voids, low tolerance of poorly mating surfaces, and high processing temperatures (roughly 2/3 rds of the parent material's melting point). As a result, diffusion bonding is mainly limited to steel (and titanium) laminations [17].Brazing involves introducing molten filler metal (e.g., copper, silver, nickel) between laminations to wet the mating surfaces of the joint and formation of a meturallurgical bond.

Bonding is achieved by diffusion of214Worked Materialsheet steel or aluminumplastics *aluminumplastics **plastics ***plastics +composites ++the filler material into the parent material and recrystallization of the filler material around the asperities thereby creating an enhanced surface area adhesion. Steel on steel requires a flux (i.e., non-metallic chemical compound) to clean the surface of oxides and other contaminants, facilitate easy wetting of the surface and then evaporates. Brazing is done at 450°C and above, and therefore is not recommended for relatively low melting point metals (e.g., aluminum). Advantages of brazing over other bonding methods include minimal effect on the base material composition, large bonding area for evenly distributing stresses, higher tolerance for mismatched surfaces than diffusion bonding because voids and gaps are filled. Disadvantages include the need for the joints to have controlled gaps or clearance, and fluxing.

Disadvantages include the need for the joints to have controlled gaps or clearance, and fluxing. Brazing is mainly limited to steel.Bonding with adhesives involves applying the adhesive material between the surfaces and letting it cure.Soldering and edge welding are not being considered as suitable bonding methods for the following reasons:Soldering is similar to brazing except that wetting of the filler material (e.g., tin, lead and their eutectic alloys) on the base material occurs at below 450°C. The measured strength is comparable to adhesive bonding but the processing is more difficult because gaps or clearance is needed between laminations. Initial experiments have shown that using soldering to bond the aluminum pieces together with minimal clearance led to poor bonding.

Initial experiments have shown that using soldering to bond the aluminum pieces together with minimal clearance led to poor bonding. Messler [16] suggests that adhesives would be a better choice than soldering because gaps won't need to be maintained with the former.Welding of the lamination edges is not being considered because the piece-wise continuous surface afforded by the beveled edges would be significantly altered and secondary machining operations would be required as evidenced by Azuma [7], Pridham and Thomson [8] and Kunieda et al. [12]..Experimental ProceduresVarious experiments being performed will establish the important mechanical and physical properties of various types of bonded lamination joints. Since the joints will add some degree of elasticity to the die, tensile and shear elastic moduli are important for the die designer to know. The types of laminated die failures that can occur due to the manufacturing processes listed in Table 1 are tensile (Fig.

The types of laminated die failures that can occur due to the manufacturing processes listed in Table 1 are tensile (Fig. 2a), shear (Fig. 2b), and peeling (Fig. 2c), either separately or in some combination. For this reason, tensile, shear and peel strengths are also useful to know. Since the bonded interfaces provide a resistance to thermal conduction in the direction perpendicular to the bond plane, as shown in Figure 2d, the thermal contact resistance for a particular type of bond is also important. The degree of fluid (i.e., hydraulic oil or air) permeation in the lamination bonds is important to know when conformal cooling with pressurized fluid (e.g., injection molding), as shown in Figure 2e, or drawing of a vacuum (e.g., thermoforming) is being considered. For this reason, the specific permeability for various bonds is being determined.

The basic premise of this test is to pull (using a Universal Testing Machine) a bonded specimen of area (A bond) and thickness (t), as shown in Figure 3a, perpendicular the bonded surfaces and measure the elongation (6.!ailure) and force (Fjailure) of the joint at failure with respect with its original position and unloaded state. The tensile forces applied must be uni-axial in order to achieve pure tension. This test is based on ASTM Designation: D 2095-72 standard test method. Assuming that the bond material obeys Hooke's law up to failure, then the tensile strength and elastic modulus are calculated using the following equations:a = tFtfailure and E = - - - - - C;failure • Abond Abondt(1)Shear Strength (Ts in MPa) and Elastic Modulus (Gs in MPa) - The test being used to measure the shear strength and elastic tensile modulus of a lamination joint is a modification of ASTM Designation: D 2295 - 92 standard test method.

The basic premise of this test is to pull a lapped joint of thickness (t), as shown in Figure 3b, parallel to the bonded surfaces and measure the elongation (6.!ailure,s) and force (Fjailure,s) of the joint at failure with respect to its original position and unloaded state. Assuming that the bond material obeys Hooke's law up to failure, then the shear strength and elastic modulus are calculated using the following equations:Flailu(~lailur, = Fand G, ="" AbondA . bond1tant The small angle approximation for the shearing strain (i.e., r = Gs because 6.!ailure.s is typically much larger than t.failure,sJ(2)) is not used in determiningPeel Strength (ap in ~) - The T-peeltest for determining the peel strength of alamination joint of width (w) is based on ASTM Designation: D 1876-93 standard test method. Peel strength is a measure of a joints resistance to localized stresses and failure by progressively opening the adhesive face normal to its bond line.

Peel strength is a measure of a joints resistance to localized stresses and failure by progressively opening the adhesive face normal to its bond line. The basic premise of this test is to pull a T-216joint, as shown in Figure 3c, normal to and at one edge of the bonded surfaces and measure the Fma~,p;Fmin.p, of the joint at failure with respect to its unloaded state.average force, i.e., Fave,pPeel strength is calculated using the following equation:a p =F/ailure,pw(3)w=25 mm(a)Figure 3 - Specimens for measuring joint strength and elastic modulus when subjected to (a) tensile, (b) shearing and (c) peeling loads.To quantify the effect of elevated temperatures (particularly the maximum values listed in Table 1) on bond strengths and elastic modulus, the tests will be run at room temperature (i.e., 24°C), 100°C, and 200°C and 1000°C (diffusion-bonded steel only).

As shown in Figure 4, uniform heating of the specimens is achieved by wrapping resistive-type heating tape around the bonded area. Using heating tape is an inexpensive means for heating the localized area up to 200°C without interfering with the failure mode of the specimen. The temperature of the bonded area is monitored by a digital thermometer using a K-type thermocouple in direct contact with the test specimen. This simple arrangement has been shown to maintain a uniform temperature to within ±3°C. These heated specimens were pulled apart on a Universal Testing Machine at a rate of 25 mmlmin.Figure 4 - Simple configuration for uniform heating of test specimens.The thermal contact (conductive) resistance per unit area (R;',c) of the various joints willbe measured using the apparatus shown in Figure 5a.

Because of the symmetrical nature of the217heating element sandwiched between identical bonded metal laminations, the heat rate in thepositive and negative x-directions is assumed to be the same (i.e., determining the value of this contact resistance is:). The equation for" _ [A(T; - T:) _~] Rt c - 2 m qx 'k(4)where:A = cross sectional area of bonded joint Ti = temperature at heating element Ts = temperature at outside surface qx = total heat rate of heating element L = thickness of single lamination km = thermal conductivity of lamination material.The specific permeability (k) of a bonded joint (i.e., how well it seals) will be measured using the experimental arrangement shown in Figure 5b.

Assuming that the bond has some degree of permeability, the test fluid used is incompressible, and fluid flow thropgh the bond is laminar, the specific permeability of the bond material is based on the Darcy's equationk =Qj.ih A/(pt - Po)(5)where:Q= volumetric fluid flow rate (measured) f1 = viscocity of the fluid h = width of the bond A/ = 1[(d i + h)t = total cross-sectional area of bond perpendicularto the direction of flowt = thickness of the bond Pi = internal pressure of the bonded cylinder and po = atmospheric pressure.A=cross(a)sectional areaHeating Element (qx=constant, qy=O)Figure 5 - Experimental set-ups for determined (a) thermal contact resistance and (b) specific permeability of a bonded lamination joint.218Preliminary Experimental ResultsSome preliminary results are described for adhesive bonding of aluminum laminations.

Four high strength, high-temperature adhesives (representative of the wide variety available) were tested including Devcon Plastic Steel® epoxy, Cotronics® Alumina-based Adhesive, Dymax 846® and 828® epoxy. Each adhesive consists of two parts, i.e., an activator (or hardener) and the base adhesive. The Devcon Plastic Steel adhesive is a metal-filled epoxy which cures to full strength at room temperatures in 24 hours, or at 120°C in 15 min and cooled back down to room temperature in 1 hour. The Cotronics adhesive system is a ceramic based adhesive that uses alumina oxide as a base adhesive. The alumina bonds to the oxidizing layer on the aluminum and creates the bond. The alumina based adhesive cures at room temperature for a 24 hour period or at 120°C for 2 hours. With the alumina based system, fast heating will create bubbles within the adhesive joint and must therefore be avoided.

With the alumina based system, fast heating will create bubbles within the adhesive joint and must therefore be avoided. Dymax 828 and 846 adhesives are polyurethane oligomer-based mixtures that can cure at room temperature in 12 hours or at 100°C in 10 minutes. Devcon is widely-available and low cost, Cotronics gives a high temperature performance, and Dymax gives the lower viscosity and is the most elastic of the group, achieving elastic shear moduli of9.1 and 7.5 MPa.The tests show that there is a correlation between the distribution of the adhesive on the surface and the bond strength. A good distribution of an adhesive is when the adhesive has spread evenly and no voids are present. According to test data, the better and more even the distribution, the greater the failure strength. All samples broke in either a full cohesive failure the (see Fig. 6a) or a partial cohesive failure (see Fig. 6b), suggesting strong bonds at adhesive/adherend interface.

All samples broke in either a full cohesive failure the (see Fig. 6a) or a partial cohesive failure (see Fig. 6b), suggesting strong bonds at adhesive/adherend interface. Full cohesive failure mode suggests that the adhesive was used to the fullest, while partial cohesive failure suggests one of the following; sample surface was the surfaces were not completely free of contaminants), gap spacing under-prepared (i.e., between surfaces was uneven (attributing to thicker adhesive sections), or contaminants were introduced during processing (or heating).The shearing experiments for every adhesive have shown that the slope of the shear force versus bond elongation curve is basically linear, i.e., obey Hooke's law, up until the point of failure. This implies that Equation 1 is valid for adhesive bonding. The shearing force and elastic moduli are calculated using Equ.

(Oq24 24 24 24 100 100 100 100Adhesive thickness (mm) 0.19 0.18 0.23 0.47 0.14 0.24 0.21 0.29Failure Load (kN)13.0 17.5 14.1 2.47 1.65 10.4 9.61 2.78Shear Strength (MPa) 10.1 13.6 10.9 1.91 1.28 8.06 7.45 2.16Elongation at Failure (mm)2.1 2.4 2.1 0.38 0.25 1.9 1.8 0.53Shear Modulus (MPa) 6.8 9.1 7.5 2.8 1.2 5.6 5.1 2.0Cohesive Failure ModePartial Full Full Partial Partial Full Full PartialFuture WorkThe design and performance issues have been identified for laminated tooling and experimental set-ups have been devised for quantifying pertinent physical and mechanical r"s, Gs, and ap experiments, aluminum laminations will be tested at 25, properties. For the at, Et, 100 and 200°C. In addition to measurements at these temperatures, diffusion-bonded and brazed steel will be tested at 500°C which is the nominal operating temperature for metal extrusion. R;',cand k will be measured at 25, 100 and 200°C for both steel and aluminum.

R;',cand k will be measured at 25, 100 and 200°C for both steel and aluminum. As a comparison with bonded laminations, R;' c , and k will be measured for laminations that are clamped together underuniform pressure. Experimental results will be compared with theoretical models (e.g., Stefan's equation for predicting adhesive bond strength) whenever possible.Following this experimental work, several examples of how the data can be used in designing laminated tooling for each of the aforementioned manufacturing processes will be developed. In addition, several techniques for automating the lamination bonding process (e.g., dip brazing, applying adhesive with a stationary roller) are being devised and will be demonstrated.AcknowledgementsThis work is being supported in part by the SME Education Foundation Grant #597-2423and by Rensselaer Polytechnic Institute.Bibliography[1] Walczyk, D.F.

and Hardt, D.E., "A New Rapid Tooling Method for Sheet Metal Forming Dies," Proceedings of the Fifth International Conference on Rapid Prototyping, Dayton, Ohio, June 12-15, 1994, pp. 275-289.[2] Nakagawa, T., "Rapid Prototyping Techniques in Japan," Proceedings ofthe 4th European Conference on Rapid Prototyping, Paris, France, October 4-5, 1995.[3] Soar, R.C., Authur, A. and Dickens, P.M., "Processing and application of rapid prototyped laminate production tooling," Proceedings of the 2nd National Conference on Developments in RP&T, Buckinghamshire College, England, Nov. 18-19, 1996, pp. 65-76.[4] Hart, F.V., "Mold and Mold Making Method," Us. Patent #2274060, Issued Oct. 31, 1938.220[5] Dickens, P., Simon, D.

18-19, 1996, pp. 65-76.[4] Hart, F.V., "Mold and Mold Making Method," Us. Patent #2274060, Issued Oct. 31, 1938.220[5] Dickens, P., Simon, D. and Sketch, R., "Laminated Tooling for Moulding Polyurethane Parts," Proceedings of the SME Conference on Rapid Prototyping and Manufacturing, Dearborn, MI, April 22-25, 1996.[6] Glozer, G.R., Brevick, J.R., "Laminate tooling for injection moulding," Proceedings of the Institution of Mechanical Engineers, Part B: Journal ofEngineering Manufacture, Vol. 207, pp. 9-15.[7] Azuma, K., "Manufacture of Press Die," Japanese Patent #59-91123, Issued December 8,1984.[8] Pridham. M.S. and Thomson, G., "Part Fabrication Using Laser Machining and Welding," Proceedings of the Fourth Solid Freeform Fabrication Conference, Austin, TX, August 1993, pp. 74-80.[9] Weaver, W.R., "Process for the Manufacture of Laminated Tooling," Us.

74-80.[9] Weaver, W.R., "Process for the Manufacture of Laminated Tooling," Us. Patent #5031483, Issued July 16, 1991.[10] Feygin, M., "Apparatus and Method for Forming an Integral Object From Laminations," Us. Patent #5354414, Issued October 11, 1994.[11] Jahn, E.M., "Method of Manufacturing Laminated Plastic Tooling and Tooling Produced Thereby," Us. Patent #5247861, Issued Sept. 28, 1993.[12] Kunieda, M. and Nakagawa, T., "Manufacturing of Laminated Deep Drawing Dies by Laser Beam Cutting," Proceedings of the r t International Conference on Technology of Plasticity, Vol. 1, Tokyo, Japan, 1984, pp. 520 524.[13] Nakagawa, T., Kunieda, M. and Sheng-Dong L., "Laser Cut Sheet Laminated Forming Dies by Diffusion Bonding," Proceedings of the 25th International Machine Tool Design and Research Conference, Dept. of Mechanical Engineering, University of Birmingham, England, April 22-24, 1985, pp. 505-510.[14] Engler, I., Schubert, E.

of Mechanical Engineering, University of Birmingham, England, April 22-24, 1985, pp. 505-510.[14] Engler, I., Schubert, E. and Sepold, G., "Direct Metal Prototyping with the LASP Technology," Proceedings of the International Conference NNS (Near Net Shape) '97, Bremen, Germany, April 14-16, 1997, pp. 335-341.[15] Berry, H.D., "Intricate Shapes Built Up From Stacked Stampings," Iron Age, Vol. 198, No. 26, Dec. 29, 1966, pp.44-45.[16] Messler, R.W., Joining ofAdvanced Materials, Butterworth-Heinemann, Stoneham, Massachusetts, 1993.[17] Hiatt, W.R., "Diffusion Bonding of AISI 4118 and 52100 Steels," Masters Thesis, Dept. of Materials Engineering, Rensselaer Polytechnic Institute, 1971.221222.

viewed PaperBoosting artificial intelligence in design processes by the use of additive manufacturing M. Ott*, N. Meihöfener*, R. Koch*Direct Manufacturing Research Center (DMRC) and Group of Computer Application and Integration in Design and Planning (C.I.K.), Paderborn University, GermanyAbstractAdditive manufacturing offers the option of converting digital prototypes into real structures as quickly as possible by the special property of tool-free manufacturing. However, this process can only be used at optimum speed if bottlenecks can be effectively avoided. One of these constraints is the design process. Although modern CAD systems allow a significant increase in many areas, this always requires a person with specific skills (e.g. engineer). In the field of AM in particular, more and more powerful software solutions have recently been published which accelerate the Design for Additive Manufacturing, including most CAD-tasks.

engineer). In the field of AM in particular, more and more powerful software solutions have recently been published which accelerate the Design for Additive Manufacturing, including most CAD-tasks. In many areas, therefore, attempts are already made to automate relevant design steps as much as possible, more and more using neural networks and artificial intelligence. This paper presents how and why such techniques can be used to generate three-dimensional structures quickly and efficiently in cases of deep generative design tasks.Introduction In today's product development process, technical solutions for data management and simplified geometry generation are receiving a great deal of attention, especially in the area of customization. Shorter product development cycles with a high number of variants and at the same time low unit costs are a primary goal demanded by the market.

Shorter product development cycles with a high number of variants and at the same time low unit costs are a primary goal demanded by the market. In the context of globalization, this already starts with the preparation of an offer, which is not possible without a conception of the product and therefore must be realized in a short time without much effort. Likewise knowledge generated in the enterprise is to be received and passed on. [BEEF+97, p. 65ff.] A high level of disertification requires very small quantities, which is why additive manufacturing is a very good choice for economical production due to its tool-free properties. However, processes must be optimized and automated in order to keep development costs as low as possible for unit numbers of 1.Fundamentals in neural networks One of the very important topics within this thesis will be the application of different neural networks.

Therefore, in this section, the basic structure of such a network and its main tasks will be briefly discussed. Basically, the intention of a neural network is based on the fact that it can act like neurons located in the human body [MAY96]. In an input layer all processed data are transferred to the network. These are further processed there by different activation functions of different neurons and afterwards in an output layer again delivered to the environment (s. figure 1). In the two boundary layers the human has the possibility to interact with the network, whereas in the inner layer different hidden layers are not accessible [MAY96]. One main task of such superstructures is to classify data, for example. These can be based on different bases.

One main task of such superstructures is to classify data, for example. These can be based on different bases. Both 2D image data and thus an image recognition can be possible, often this is then realized over so-called Convolutional Neural Networks (CNN), and different semantic connections among other things in Natural Language Processing applications can be a focus of neural networks. For example, the mentioned CNNs try to generate a classification from an overall image (s. Figure 1) by iteratively extracting subregions of an image [SHNA15], whereas Graph Neural Networks (GNN) link696information by the logical relation of points and edges between each other and thus draw corresponding conclusions [ZHCU20]. A special case are the so-called Generative Adversarial Networks (GAN), which generate data themselves. For this purpose, two competing networks operate simultaneously. One is the generator network and the other is the discriminator network.

For this purpose, two competing networks operate simultaneously. One is the generator network and the other is the discriminator network. The generator network tries to generate real-looking data, whereas the discriminator tries to recognize it and declare it as fake. In this process, both networks learn from each other, so that after a large number of iterations, very real-looking data is produced. However, care should be taken to ensure a good balance between the two networks, as otherwise it favors a very unstable relationship [LULI21].0.95 dog0.05 Non-dogInputHidden LayerOutputFigure 1: Schematic representation of a neural network for classificationFundamentels in Product-Life-Cycle Within the product life cycle, a product passes through various phases.

A rough division can be made here as follows:Productdevelopment • Production and • After-Sales.Product development is characterized by planning, design and production preparation before the component is finally physically manufactured. Subsequently, further support tasks arise, which appear more topical than ever, especially with regard to sustainability aspects. [SZBE07] Figure 2 shows the course of the product life cycle in simplified form.ProductionProduct-planningProductdevelopmentUsageProductionProduct-Life-CycleConstructionProduction-preperationAfter-SalesRecyclingFigure 2: Course of the Product-Life-Cycle in simplified form [SZBE07]697In the context of this work the range of the product development stands above all in the foreground. Already in the product planning and construction a large part of the costs of a component are specified and the possibilities of influencing these for potential changes decrease with progress of the process.

The well-known Rule-of-Ten shows this in Figure 3.0.1101001Construction and DevelopmentProcurement and ProductionDelivery and UseCosts per mistakeFigure 3: Rule of Ten [VDI2247]In earlier processes, the first phases of product development in particular were characterized by a great deal of creative work. The use of Computer Aided Design (CAD) systems first appeared in later processes, the actual design, which is why the general use of software in product planning did not find much application. The following figure shows that the development effort has a relatively large margin in the transition area between the processing by design methodologies (the creative work in product planning and conception) and the elaboration in design (the resource-intensive work).

Since a primary goal is to keep development times short and thus reduce the development costs incurred, it is advisable to make further use of this space [SKBL06].Concept100%85%70%50%30%10%0% 0 - 10%50 - 60%20 – 40%ca. 15% othersDevelopmenttimeDevelopmenteffortsWork through construction methodsWork through CAD and construction methodsProductionCreative WorkPossible computational workDraftTaskElaborationFigure 4: Developmentefforts and –time during the design process [SKBL06]698Automatization in design processes and modeling Especially due to the design freedom of additive manufacturing, it is necessary to integrate and consider the corresponding knowledge about this at an early stage in the product development process. Modern automation of design processes is already attempting to enable "Design for X" in many areas, albeit with more or less success. One of the much-used software tools is Generative Design.

Modern automation of design processes is already attempting to enable "Design for X" in many areas, albeit with more or less success. One of the much-used software tools is Generative Design. In contrast to conventional topology optimization, here various input parameters of an optimization are varied in such a way that, subsequently, the user can select an optimum result from a number of design proposals on the basis of various criteria (construction time, amount of support material, thermal distortion, etc.). In addition to the topology optimization approach, there are also parameter and shape optimization approaches within structural optimization. These differ mainly with respect to three criteria. The detail quality, the computing power and the design freedom. From this field especially the parameter optimization is a much used possibility to generate semi-automated processes.

The detail quality, the computing power and the design freedom. From this field especially the parameter optimization is a much used possibility to generate semi-automated processes. Here the constructions are built up first parametrically, in order to interact then over suitable interfaces with the designer, in order to change these [BLFI05] [REI19]. Since parametric designs usually lead to conventional manufacturing methods due to their regular progressions, the greater focus here is placed on the possibilities of automated topology optimization or generative design methods. According to [REI19], there is a great need for action in the area of digital geometry processing, especially with regard to the benefits of geometries for additive manufacturing, where modern artificial intelligence systems can provide a remedy. For this purpose, however, it is first important to know in which way geometries can be modeled and how to continue working with different methods.

For this purpose, however, it is first important to know in which way geometries can be modeled and how to continue working with different methods. The following figure shows the basic classification of modeling processes according to [REI19]. In conventional CAD, the curve-based modeling processes specified there are primarily used. They make it possible to define analytically describable geometries, which the other two groups of polygons and digital sculpting cannot do. Points and surface meshes consisting of polygons can only approximate corresponding geometries [ZWE20]. The same applies to the group of voxel models.

Points and surface meshes consisting of polygons can only approximate corresponding geometries [ZWE20]. The same applies to the group of voxel models. However, since each element of voxels as well as points stands for itself and neural networks exchange and calculate their information viaModeling ProcessPolygon-basedCurve-basedDigital SculptingPointcloudsSurfaces through curvesDisplacement (SubD)Surface-MeshNURBSVolumetric(Voxel)SplinesDynamic Tesselation(Combination)Geometric primitives Figure 5: Modeling processes for geometries [REKO15]699vectors, tensors and matrices, these can be used very well to build corresponding arrays, which can function as input values. However, there are already first geometry recognition networks, which try to function on the known splines, but here the preparation of the data turns out to be much more complex [FELE18].

However, there are already first geometry recognition networks, which try to function on the known splines, but here the preparation of the data turns out to be much more complex [FELE18]. In the context of the paper a further intensification of the structure of the data contents within a file format can be renounced, since here the possibilities and application times are crucial.Artificial intelligence in design processes Artificial intelligence can be introduced at various points in the design process. For this purpose, it can be useful to link different neural networks in order to take over different tasks in the process. A very good example of how neural networks can work together to solve highly complex tasks with different characteristics is the topic of autonomous driving, for example at Tesla. There, 48 networks act simultaneously to perform all the tasks that enable driving without danger [KAR20].

There, 48 networks act simultaneously to perform all the tasks that enable driving without danger [KAR20]. In the design, different data types are present than in autonomous driving, however, depending on the time of application, a wide variety of activities are also to be performed by a neural network. In the following, the most important use cases will be shown in more detail, potential networks will be presented and a possible output will be shown.Basically, it has been shown that the product development process can be uniformly divided into different phases and that especially mistakes made at the beginning are very expensive to fix. The initial situation in the design process is therefore always a problem, whether it is based on the market or on an existing product. This then raises the question of a new or variant design. Artificial intelligence can be used and supported at various points.

This then raises the question of a new or variant design. Artificial intelligence can be used and supported at various points. In order to provide efficient support, the goal of the application must first be defined. By analyzing different scenarios, the inspiration for new designs can be defined as a clear goal in the area of new design. In variant or improvement design, on the other hand, problems that have already been identified must be eliminated or improved (s. Figure 6).This paper is initially limited to the methods of the inspiration process and the prediction of mechanical behavior, which are based on the methods of polygons and digital sculpting. Therefore,New DesignProblem / Specifications / market requirementsMech.

Therefore,New DesignProblem / Specifications / market requirementsMech. behavior predictionDetailed constructionParameter-optimizationLocation for AI-SupportInspirational processesSolved Problem -> new geometryVariant constructionFigure 6: Possibilities on AI Support in design processes the first step is to take a closer look at the inspiration process. Due to the early stage of development, data here often cannot be interpreted as finished volume models and a rather rudimentary form of representation is usually sufficient. In the context of this paper, therefore, a generative network is used, which initially has the task of generating different variants of one and the same product. This700means that the rough framework is initially clearly defined, but can differ in fine features. An example is a chair, which initially consists of a seat, the backrest and the legs of the chair. The exact detailing of these features will then be varied.

An example is a chair, which initially consists of a seat, the backrest and the legs of the chair. The exact detailing of these features will then be varied. This is precisely the functionality offered by DECOR-GAN, which is therefore used for this application. However, since a generative network can only classify between real and non-real structures, an additional classification network is needed. Since the GAN is based on voxel data and was initially trained for the data set of the ModelNet10 , the VoxelNet was used for the classification of the resulting geometries, which offers a very good cost-benefit ratio.1) Pure Inspiration ProcessIn the state of the art, the various phases in the product development process have already been addressed. Here in particular, it can be seen that the early phases are characterized by creative work, where the pure CAD benefit is rather low.

Here in particular, it can be seen that the early phases are characterized by creative work, where the pure CAD benefit is rather low. In order to promote further ideas here and to search for possibilities for a potential design, supporting methods of artificial intelligence can be used here. Since it is not yet necessary to generate complete geometries that can be described analytically, it makes sense to use a polygon-based or voxel-based model. In the area of polygons, point clouds would be suitable here, since these cannot initially have any errors, since each point stands individually for itself. In the image processing area, we would be talking about pixels here. Voxel models offer the same properties here, but in three-dimensional space, and are therefore often referred to as 3D pixels. But here too, the absence of a single voxel exorbitantly disturbs the overall volume or representation.

But here too, the absence of a single voxel exorbitantly disturbs the overall volume or representation. In the case of surface models, on the other hand, the absence of a surface leads to the fact that it is no longer possible to distinguish between inside and outside and thus no volume can be defined [REI19].In the pure inspiration process, it is first important to differentiate which tasks the neural network should take over. It has already been shown that the VoxelNet (the network can be found at: https://github.com/dimatura/voxnet) provides a good tool to recognize geometries on a voxel basis. Training was done here with the ModelNet 10 / 40 dataset ( https://modelnet.cs.princeton.edu), which is a freely available dataset with 10 and 40 classes, respectively. During the research, the network was used for two different types of classifications. This again depends on the data preparation which is exemplary shown in Figure 7.

During the research, the network was used for two different types of classifications. This again depends on the data preparation which is exemplary shown in Figure 7. Since the training data requires a uniform input vector, the data must first be normalized for this purpose. Depending on the use case, two possibilities can be identified here. If only the geometries have to be recognized as such, for example that an airplane is an airplane, then it is sufficient to define a voxel space and to scale all geometries with their maximum dimension into it. However, for small classes of the data set, the details or features that precisely describe an object are omitted.

However, for small classes of the data set, the details or features that precisely describe an object are omitted. If these are to be recognizedPredictionN-dimensional data for trainingVoxelNet (3D CNN)Training Geometry (.step, .x_b, .stl)LabelValidation Geometry (.step, .x_b, .stl)NormalizationNormalizationN-dimensional data for trainingVoxilizationVoxilizationFigure 7: Data preparation for the VoxelNet use701however it offers itself to define the voxel space size in dependence of the geometry size. This makes it possible to recognize, for example, a two-engine or four-engine airplane or other features. Further input parameters are the labels of the individual data sets. Depending on the resolution of the voxel space, different sized batchsizes can then be chosen to train the mesh. We chose a 32x32x32 voxel space and a batch size of also 32.

Depending on the resolution of the voxel space, different sized batchsizes can then be chosen to train the mesh. We chose a 32x32x32 voxel space and a batch size of also 32. With this we achieved an accuracy in ModelNet 10 of 87% without any data augmentation, which is a very good reference value [MASC15]. In order to be able to use this principle for inspiration, further steps must be added. For this purpose, geometries must be generated automatically on the basis of certain properties or labels. This can be done very well with the "Generative Adversarial Networks (GAN)" explained in the basics. As already described above, these are able to generate geometries which are then classified as artificial or real by the network itself. Here, too, various networks have been selected within the scope of the work on the topic in order to be able to generate corresponding geometries automatically.

Here, too, various networks have been selected within the scope of the work on the topic in order to be able to generate corresponding geometries automatically. Based on the classification of the modeling processes, care has been taken to include the two classes of polygon-based processes and digital sculpturing. The data preparation is to be equated here first with the procedure of the VoxelNet described before. Here it is to be paid attention only accordingly whether on voxel-basis or on point-basis one works. Accordingly, if necessary, the point of "voxilizing" must be replaced by "generating point clouds". In the voxel area, the DECOR-GAN (code available at: https://github.com/czq142857/DECOR-GAN) has been used here. This offers especially the possibility to generate exact features in geometrically similar elements.

This offers especially the possibility to generate exact features in geometrically similar elements. For example,Figure 8: Possible chair inspiration generated by a GAN on voxels702if the class of the chair is considered from the dataset of the model net, a chair usually has some legs, a seat and a back. How exactly these are designed, however, is now to be determined in the design process. Various creativity methods can be used to remedy this situation. Alternatively, GANs can be used to accelerate these tasks. The figure 8 shows voxel-based chairs that have been classified as real objects by the network.2) Prediction of mechanically stressed regionsA third neural network, which will be presented in this paper and has been applied in the work, is the DiffusionNet (code available at: https://github.com/nmwsharp/diffusion-net).

It is based on the principle of natural distortion and manages to achieve extremely high accuracy in classification and segmentation independent of whether point clouds or polygon representations are used. Another positive aspect for the use of the network is the small number of samples needed to train because of its natural diffusion. In combination these two factors made the decision for using this network compared to others. In this paper, the segmentation was abstracted to generate stress predictions in components. For this purpose, the occurring stresses were first divided into classes and processed as a training set. With an accuracy of slightly over 93%, the stress predictions here are at a very good level for identifying any excess stresses. The training set here consisted of various simulations and stress values of a bracket optimized for additive manufacturing.

The training set here consisted of various simulations and stress values of a bracket optimized for additive manufacturing. In the figure, the left side shows the result from the simulation, while the right side shows the prediction of the mesh. In the prediction, even smaller edges can now be seen in the colored segmentation, which is due to the division of the stress values into certain classes (stress ranges). The more classes are defined here, the more exact the colored course becomes.Figure 9: Left: Stress result from FEM ; Right: Stress Prediction from neural networkImpact and Summary Along the elaboration, the main topics of additive manufacturing, time saving in development and the topics of artificial intelligence were addressed again and again. Especially the first partially automated possibilities of Generative Design try to make these goals usable with as little user experience as possible and extremely fast computational algorithms.

Especially the first partially automated possibilities of Generative Design try to make these goals usable with as little user experience as possible and extremely fast computational algorithms. However, similar to conventional topology optimizations, models must be built and calculated here as well, which in turn limits the time savings and thus the cost reduction per component. If the three neural networks presented and their functions highlighted in the paper are now considered in the context of the703overall system, it is noticeable that there is a large overlap here if these functions can be coupled with each other. At this point it must be mentioned that this state is future work and right now just hypothetical.In summary, the relationship described at the beginning of this paper between additive manufacturing and the benefits of artificial intelligence in the product development process can be assumed to be a direct causal relationship.

It has been shown that the early phases of product development are characterized by work that has so far received little support from modern technologies, but which therefore requires a great deal of time, especially for small quantities, and this leads to high development costs. Individual solutions were shown, which can already be implemented with today's methods and that in combination of these possibilities there is a much greater potential, which must now be explored. Here, too, a potential solution approach for the further and profitable use of neural networks in the design process has been created, which should be used further, especially in combination with additive manufacturing.

ICK PRINTING: FREEFORM FABRICATION OF MODULAR ARCHITECTUAL ELEMENTS WITH EMBEDDED SYSTEMS Jeffrey I. Lipton1, Daniel Cohen2, Hod Lipson2 1Applied and Engineering Physics 2School of Mechanical & Aerospace Engineering, Cornell University, Ithaca, NY 14850 USAAbstract We propose the use of modular, printed bricks to enable the integration of building systems and various processing techniques through the use of scalable printer platforms. This is enabled by a novel material platform comprised of clay, gypsum cement, FabEpoxyTM, and SS-26F conductive silicone. On an open-architecture SFF system, a segment of cement wall with embedded electrical and fluidic conduits and various processing techniques was fabricated. Electrical and fluidic tolerances were comparable to traditionally constructed systems.IntroductionTraditional building techniques tend to create structures with right angles and vertical walls.

Electrical and fluidic tolerances were comparable to traditionally constructed systems.IntroductionTraditional building techniques tend to create structures with right angles and vertical walls. The majority of buildings are made from discrete parts such as I-beams, bricks, or wood beams. These discrete building elements naturally lend themselves to the creation of simple Cartesian buildings. While bricks have been used or centuries to create curved and complicated structures, the surface retains a certain granularity, and the process is labor intensive and requires onsite modification of the bricks. There are several techniques for homogenous structure creation. Primarily, these focus on casting concrete in various ways. Insulating concrete forms allow for the creation of structures by creating a casing for the structure using preformed parts. Spaces are left for the installations of electrical and plumbing work after the concrete has cured.

Spaces are left for the installations of electrical and plumbing work after the concrete has cured. With shotcrete, a wire mesh or inflated air form provides the shape of a structure which is covered in the concrete. These more naturally lend themselves to curved structures, but lack the resolution that SFF can offer.SFF presents a solution to this problem by enabling the creation of arbitrary three- dimensional structures. The Radiolaria pavilion, designed by Shiro Studio and D-Shape, is a building being printed on the world’s largest SFF system. It is constructed from 5-10 mm layers of artificial limestone (1). The Contour Crafter, developed by Behrokh Khoshnevis, at the University of Southern California, relies on the wet extrusion of cement and other materials to from a square tube and smoothing via a mechanically actuated trowel (2). Both systems create a structure of homogenous and continuous composition.

Both systems create a structure of homogenous and continuous composition. They are incapable of using multiple material processing and printing techniques. While structures of any form could be created, the building systems would be bound by their prefabricated geometry. Any structure printed would be bound to the geometric limitations of its building systems. Additionally the printers for both systems are designed to encompass the built structure. The speed, size and resolution of the printed structure would be limited to that of a single printer – in that sense, the system does not scale favorably with the size of the structure.A structure made from discrete elements would have the ability to take on any geometry with the added advantage of being produced from many lower cost machines in parallel.

The speed of construction would be limited only by the number of printers, and printers could be specialized (at the sacrifice of interchangeability) to create higher resolution building blocks where needed. Lawrence Sass’s work with digital home fabrication follows similar principles. His structures are created from laser cut plywood. The laser cutter is significantly smaller than the structure created, and generates structurally sound buildings (3).All current building fabrication techniques require the installation of building systems, such as piping and electrical work, after the structure has been completed. This separation can lead to construction delays since a wall cannot be completed until all requisite electrical and plumbing work is complete. Often the workers who install the structure are separate from those who install plumbing and wiring.

Often the workers who install the structure are separate from those who install plumbing and wiring. The novel printing materials platform developed could enable SFF to successfully co-install electrical and plumbing infrastructure in arbitrary geometric structures. This could enable a new and uncharted architectural design space. (Figure 1)Figure 1: Illustration of Custom geometry, printed, interlocking bricks used to construct a structure larger than the printer. The interference fit of the bricks locks them into position.BackgroundThe Vast majority of bricks created in the world today are fabricated using extrusion processes where clay is forced into a die and then cut into desired lengths. This mass manufacturingtechnique is prevalent in the industrialized world. Bricks generated using this system usually have excellent mechanical properties, but vary only in length and cannot take on arbitrary shape without the retooling of the assembly line.

Bricks generated using this system usually have excellent mechanical properties, but vary only in length and cannot take on arbitrary shape without the retooling of the assembly line. The bricks are fired in a rail kiln at temperature ranging from 1000 to 1500 degrees depending on the clays composition. (4) In the more rural settings, brick making is a manual process. Clay is placed by hand into preformed dies. These bricks are then placed in static kilns or Bull’s trench kilns. (5) The firing process of both methods is shape independent and would not be affected by the use of SFF’ed bricks.In order to demonstrate techniques and fabricate with the new materials platform in this paper, a Model 1 Fab@Home printer was used. Fab@Home is an open source, low-cost SFF system. The printer automatically moves the syringe and dispenses material in a process known as “robocasting”.

Fab@Home is an open source, low-cost SFF system. The printer automatically moves the syringe and dispenses material in a process known as “robocasting”. Its small reservoir and size limits that scale of prints, however its design allows for the combination of any extrudable materials into a printed object.Figure 2: A Model 1 Fab@Home 3D printer with various printed objectsMaterialsHydro-Stonetm Super-X gypsum cement from United States Gypsum was selected as printable cement since it has a fine powder size and high strength, and does not separate under pressure. However its self leveling nature prevents its use in printing without a casing. Xanthium gum was used as a viscosity modifying additive that allowed the cement to retain a given shape while it cures. Xanthium gum was selected because small concentration of the material drastically changes the viscosity of water-based solution.

Xanthium gum was selected because small concentration of the material drastically changes the viscosity of water-based solution. The shape-retaining cement mixture is prepared by adding 0.5 parts Xanthium gum powder by weight to 100 parts hydro-stone super-x powder by weight.The powders are mixed thoroughly to give an even distribution of materials. This was accomplished by high amplitude low frequency vibration or mechanical agitation. In order to create cement the mixture must be added to 24 to 30 parts water for every 100 parts mixture. Higher water content produces a less dense and lower viscosity material. Hydro-Stone Super X cement is used with 21 to 23 parts water to 100 parts powder (6). The powder must be added to the water slowly while the water is stirred in order to ensure the production of a homogenous material.

The powder must be added to the water slowly while the water is stirred in order to ensure the production of a homogenous material. The mixed cement has a work life of 18 to 25 minutes depending on the concentration of water and uniformity of material.FabEpoxy by Kraftmark was used as a pipe and conduit material due to its high strength and negligible shrinkage (7). The FabEpoxy used was mixed via a static mixer attached to a two- material cartridge. Once the material had been extruded through the entirety of the mixer, the static mixer was attached to a sealed 10ml Engineered Fluid Dispensing syringe. This minimized the ability of air to enter the syringe and ensured a perfect mixing ratio. Once mixed the material has a work life of 3 hours, and takes up to 24 hours to cure.

This minimized the ability of air to enter the syringe and ensured a perfect mixing ratio. Once mixed the material has a work life of 3 hours, and takes up to 24 hours to cure. The cure rate can be increase through the application of heat below 80 degrees Celsius.SS-26F a 1-part, silver-filled RTV silicone, was selected as an electric conductor because it has been well characterized for use on the Fab@Home platform. It is available in small quantities directly from the manufacturer. It does not form good contact between cured and uncured material and therefore should not be allowed to sit for more than 10 minutes while printing (8).Kentucky stone ball clay was selected for ceramic brick printing due to is fine powder size. The moisture level of the clay dramatically affects performance of a print and must be controlled. The material is prepared by mixing 70 parts clay powder to 30 parts water, by weight.

The moisture level of the clay dramatically affects performance of a print and must be controlled. The material is prepared by mixing 70 parts clay powder to 30 parts water, by weight. The water saturated clay must be thoroughly mixed to prevent the formation of non-uniformities. The mixed clay must be stored in a sealed water tight container in the absence of air to prevent the clay from drying out.Printing ProcessThe fabrication was done using a two-syringe, Model 1, Fab@Home printer. Each material needs to be calibrated to give the machine a list of parameters. Primary parameters affect the path planning of the software and determine the resolution and feature size. Secondary parameters, such as “suckback” and “pushout”, affect the machines ability of produce a product that matches the geometry and ensures a good finish to the part.

Secondary parameters, such as “suckback” and “pushout”, affect the machines ability of produce a product that matches the geometry and ensures a good finish to the part. For the novel materials, a series of trial and error modifications are needed to determine the parameters (9).For materials other than FabEpoxy, the wiper piston was removed from the syringes and material was added through the top opening. Material was spread across the top to create an air tight seal. While adding material, pressure is applied on the material to force it down into the syringe. While the material is even with the top opening, the wiper is re-inserted to prevent air form being trapped between the material and the wiper. Air is trapped in the bottom of the syringe by this procedure and requires that the first cubic centimeter of material is extruded through the bottom opening before use.

Air is trapped in the bottom of the syringe by this procedure and requires that the first cubic centimeter of material is extruded through the bottom opening before use. Large air pockets affect the material’s print parameters, and small pockets create imperfections in the printed object.The minimum feature size obtainable on the Fab@Home is determined by the smallest EFD tip that can be used in extrusion. Since Fab@Home motors produce a force less than average human strength, a material must be extrudable by hand for it to be printed. Materials are placed in the syringe using a method described above. Pressure is applied after successively smaller tips are used. The last tip to successfully allow extrusion without clogging is used. If there is significant clogging during printing of homogenous materials, the next largest tip is used.While the printer deposits the clay, a humidifier moistens the surface of the printingobject to promote interlayer bonding.

Without the additional humidity, the material will extrude improperly and form regions where the material will not bond (see Figure 3).(a)(b)Figure 3: Parts printed without additional humidity(a) tend to bond improperly between layerscompared with those printed in a humid environment(b)After an object is printed using clay, it must be dehumidified. This can be accomplished by either letting the object sit under a halogen lamp for 12 hours, or heating the objects to 50 degrees Celsius for several hours. Once dehumidified the clay is heated up to 1000 degrees C at a rate of 300 degrees per hour. The clay is held at that temperature for 24 hours. After the clay has been fired, it is cooled at a rate of 300 degrees Celsius per hour until it reaches room temperature.A significant source of error in the printing of objects is uneven drying. This can be due to a non-uniform heat source or a barrier to moisture.

FabEpoxy machinable nature lends itself easily to joining a printed fluidic channel to existing plumbing infrastructure. National Pipe Thread (NPT) inserts can be screwed into a FabEpoxy tube while wrapped in Teflon tape. This creates a water and air-tight seal which is an ideal joint between traditional and SFF’ed components.(a)(b)Figure 5: A FabEpoxy pipe (b) can be joined to standard plumbing and pneumatic fixtures via NPT plumbing inserts and barb adapters (a).The Uniform Plumbing Code (UPC) calls for water to be delivered to homes for domestic use at between 50 to 70 PSI. Valves and supply lines as well as appliances are designed to withstand up to 80 psi (10). In order to determine the viability of using FabEpoxy as a fluidic conduit, a test sample was produced using an inner diameter of 12mm and a wall thickness of 6mm. The sample was allowed to cure over night under a halogen lamp, to speed curing, in order to ensure maximum strength had been achieved.

The sample was allowed to cure over night under a halogen lamp, to speed curing, in order to ensure maximum strength had been achieved. NPT inserts were used to attach the pipe to an air supply, and to seal the pipe. The pipe was allowed to pressurize up to 100 PSI and drained repeatedly. The 12mm inner diameter pipe was able to successfully survive pressurization up to 100psi, without any noticeable deformation. FabEpoxy can withstand temperatures up to 121 degrees Celsius, well above home water temperature.FabEpoxy piping was successfully embedded into a printed cement structure by inserting a preprinted pipe into a printed section of cement, and by simultaneously printing both materials.

The outer diameter of the pipe was .2mm smaller than the inter diameter of the cement in order to account for inaccuracy of the tip heads and printing errors.(a)(b)Figure 6: FabEpoxy pipes can be embedded into cement structure by simultaneously printing the materials (a) or by mechanical insertion after successful printing (b).In order to demonstrate the ability to SFF an electrical conductor inside of cement, a section of SS-26F was printed inside of a gypsum cement casing. After allowing 24 hours for curing, the sample’s resistance was measured. The sample had a resistance of 0.2 Ohms. SS-26F has a build resistivity of 5.0 x 10-6 Ωm, while copper has a resistivity of 17.2 x10^-9 Ωm (8). Therefore a printed wire would need to have a radius 17.04 times greater than an equivalent copper wire.

Therefore a printed wire would need to have a radius 17.04 times greater than an equivalent copper wire. A 14 gauge copper wire would correspond to a 13.87mm diameter printed wire.Figure 7: A section of printed electrically conductive silicone embedded in a printed cement structure.Ball clay has a notoriously high shrinkage rate that makes is a poor choice for most potter applications. In SFF this can be advantageous since it allows for the fabrication of parts which have a smaller feature size post firing. 10 samples of bricks printed with a geometry of 20mm x 20mm x 4.4mm. After firing the part shrank by 14.68% with a standard deviation of 1.47% in the plane of the layer, and by 16.33% with a standard deviation of 2.52% in the direction perpendicular to the layers of the object.

Using this as a directional scaling factor, it should be possible to account for shrinkage by rescaling an object at printing time to ensure it meets the intended geometry.An interlocking chain of bricks was fabricated using Kentucky clay. Since the clay shrinks considerably, an interlocking dovetail design was selected since it allows for limiting 2 degrees of freedom without the need for extraordinarily high geometric fidelity. Figure 5 shows a series of bricks with dovetail interlocks before firing in a kiln. The Fab@Home printer used to make the bricks has a build space of 17.8 cm x 17.8cm. The one dimensional structure has a length of 22.9 cm.(a)(b) Figure 8: Interlocking Ceramic bricks (a), next to standard ruler, formed into a structure(b) on the scale of the printer.Discussion:individually printed bricks, successfully demonstrates a printer’s ability to generate architectural structures of arbitrary scale using parts bound by the build size of the printer.

As a demonstration of the ability to fabricate decorative objects, a brick was created with the Cornell Computational Synthesis Lab logo embedded as a relief. Each shape in the tangram logo was placed at a different depth in the brick. The minimum tip size affected the fidelity of the print since it forced all corners to be rounded with a fillet of radius 0.4191mm. Using this technique it should be possible for any design to be placed onto the bricks used to make a building. Often in classical style architecture, figure heads and text are added to the structural design of the building as decoration.

Often in classical style architecture, figure heads and text are added to the structural design of the building as decoration. The technique would greatly reduce the cost of such architectural flourishes by removing any skilled labor or work by hand.The printed superstructure, comprising(a)(b)Figure 9: The CCSL logo (b) printed as a relief in a ceramic brick.In order to demonstrate the effectiveness of the printed cement with electrical conductor, a candelabra bulb socket was embedded inside a printed cement object along with wires attached to the socket. The printed wires were then connected to 120 volt AC power via a standard power cord. When the cord was plugged in, the printed wires successfully carried current to the socket and activated the bulb.

When the cord was plugged in, the printed wires successfully carried current to the socket and activated the bulb. The conductive silicones pliable nature easily lends itself to merging traditional and printed electronics since it can adapt to the unique geometry of the coupling wire or lead.(a)(b)(c)Figure 10: A light bulb attached to a socket, embedded in a cement structure(A) and powered via printed conductor before (B) and after(C) current is appliedAs a test of the feasibility of the brick printing platform, the flow rate of material wasdetermined for various drive systems. Material was loaded into a 10ml EFD syringe and placed in various deposition systems. The time to empty the container of its volume of material was record, and flow rate determined (See Table 2). A circular room with 100 square foot (9.29 m2) interior and 3 inch (7.62 cm) thick 10 ft (3.048 m) tall walls requires 90.58 cubic feet (2,565,013 cm3) of material.

A circular room with 100 square foot (9.29 m2) interior and 3 inch (7.62 cm) thick 10 ft (3.048 m) tall walls requires 90.58 cubic feet (2,565,013 cm3) of material. A current Fab@Home design would take almost 10 years to print the clay needed to build the structure, provided it worked non-stop with constant material feed. However a more advanced printer with a single, 1.524 mm tipped, tool 90psi driven tool would take 5 days and 11 hours. 6 printers with sufficient path speed would be able to print the material in less than a day using the same tools.There are tradeoffs between time and resolution and parallelization. One can sacrifice resolution for build speed. Normal construction techniques do not provide 0.8mm resolution. With printed structures, the high resolution is only needed for exteriors and interfaces with building systems. Using a dual-resolution printer, one could balance the need for resolution and print speed.

Values are for machines using 90 psi tools to deposit gypsum cement.A potential application of brick printing is the use of local materials for the construction of unique structures in remote and impoverished areas. The small size and scalable nature of the brick printing system allows for a minimal upfront capital investment. A person or organization could start with several printers, and reinvest their earnings into more machinery, directly increasing productivity. The clay based printing could use local earthen material to fabricate the structure. Kilns based on the Bull’s trench design allow for the firing of large amounts of bricks with minimal resources.Future WorkA logical next step in this research is to print and entire wall power socket and lighting fixtures into a cement superstructure. In order to accomplish this, a more rigid printable conductor would be needed.

In order to accomplish this, a more rigid printable conductor would be needed. In order to create larger superstructure using this material platform a new deposition tool would also need to be developed. While a large material reservoir would allow for the easier creation of larger models, the short work life necessitates the creation of a mix-on-demand tool that would mix the powder and water of the gypsum cement shortly before printing. This new tool would not be constrained by the pressure and tip size limitations of the current Fab@Home tool and could increase the resolution of cement printing.ConclusionSolid Freeform Fabrication has great promise in the field of building construction. While most arbitrary geometry structure fabrication techniques rely on the use of large scale printers, small size printers with high resolution could prove to be equally useful..

While most arbitrary geometry structure fabrication techniques rely on the use of large scale printers, small size printers with high resolution could prove to be equally useful. . Current large scale printers can create structural elements but cannot embed building systems during the printing process. The combination of cement structures with FabEpoxy pipes and extrudable conductor can allow for the seamless integration of building systems into the printing process. Printed bricks can be used to create structures not bounded by the size of the printed. The printed bricks could be used as both structural support and decorative flair.

Printed bricks can be used to create structures not bounded by the size of the printed. The printed bricks could be used as both structural support and decorative flair. In addition to possibly enabling new ways of constructing classical designs, this approach opens entirely new design paradigms in which designers are less constrained by the limitations that current fabrication technologies impose.AcknowledgementsThis research was supported in part by NSF Creative-IT grant 0757478 and the Defense Advanced Research Projects Agency (DARPA), Defense Sciences Office (DSO) under the Programmable Matter program.References1. D-Shape. The technology. D-Shape. [Online] March 19, 2009. [Cited: June 2, 2009.] http://www.d- shape.com/pg2.htm.2. Khoshnevis, Behrokh. AUTOMATED CONSTRUCTION BY CONTOUR CRAFTING. Journal of Automation in Construction. 2004, Vol. 13, 1.3. Instant House: Onsite manufacturing of housing with Digital Fabrication. Lawrence Sass, Marcel Botha. 04, s.l.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceBubble Formation in Additive Manufacturing of Borosilicate GlassJunjie Luo*, Theresa Bender*, Douglas A. Bristow*, Robert G. Landers*, Jonathan T. Goldstein†, Augustine M. Urbas† and Edward C. Kinzel* *Mechanical and Aerospace Engineering, Missouri University of Science and Technology, 400 W. 13th St., Rolla, MO, USA 65409; †Air Force Research Laboratory, Materials and Manufacturing Directorate, Wright-Patterson Air-Force Base, OH, USA 45433AbstractEntrapped bubbles are an important problem in conventional glass manufacturing. It is also a significant factor determining the quality of glass products produced using additive manufacturing (AM). This paper reports on the bubble formation and entrapment in filament-fed AM printing of borosilicate glass.

This paper reports on the bubble formation and entrapment in filament-fed AM printing of borosilicate glass. This process involves maintaining a local molten region using a CO2 laser. A 2 mm filament is fed continuously into the molten region while CNC stages position the workpiece relative to the laser and filament feed. Two different bubble regimes are identified in experiments with borosilicate glass. These regimes can be related to different physical phenomena, specifically, bubble entrapment at defects in the filaments and bubble formation due to reboil. These can be overcome by selecting defect free filaments and by minimizing the temperature inside the molten region to prevent breaking down the glass. Understanding these mechanisms allows bubble-free glass to be printed.

Understanding these mechanisms allows bubble-free glass to be printed. Finally, residual stress in the deposited glass pieces is removed using post-deposition annealing and validated using a polariscope.IntroductionAdditive manufacturing has been widely studied for structural parts using numerous materials, such as metals, polymers, and ceramics [1]. Several studies has considered AM of transparent optical elements. Most of these studies have focused on transparent polymer materials, including fused deposition modeling [2-4], ink-jet printing [5,6], selective laser melting (SLS) with post index matched resin infiltration [7], and multiphoton stereolithography [8].

Though these techniques are able to print transparent parts, their applications are limited since the poor material properties compared to glasses and other inorganic materials.Glasses are widely used in high quality and high power optics due to higher transmissivity, lower coefficient of thermal expansion, and more stable refractive indices [9]. However, there is limited published research on AM transparent glasses. Conventional, AM techniques have been studied for glass printing, including selective laser melting/ sintering (SLM/SLS) [10-14], and extrusion techniques [15]. However, only non-transparent structural parts have been printed using these techniques. Recent studies have demonstrated the production of 2D and 3D solid structures using transparent soda lime glass and even quartz feedstock [14-20]. Compared to soda lime glass, borosilicate glasses have a lower Coefficient of Thermal Expansion (CTE) (1/3 of soda lime glass [9]).

Compared to soda lime glass, borosilicate glasses have a lower Coefficient of Thermal Expansion (CTE) (1/3 of soda lime glass [9]). This leads to a higher thermal shock resistance that makes it a better material for applications ranging from optics to glassware. Two issues that confront additive manufacturing of glasses (including both soda lime and borosilicate) are bubble entrapment/formation [20] and residual stresses due to rapid cooling of the glass during printing.Bubble formation and entrapment is a significant issue for conventional glass industry. Bubbles cause optical scatter in glass which limits the optical performance. For example, six bubbles per ton of glass leads to a 10% rejection rate in television panel industry [21]. In addition to optical considerations, bubbles significantly weaken the glass mechanically.

For example, six bubbles per ton of glass leads to a 10% rejection rate in television panel industry [21]. In addition to optical considerations, bubbles significantly weaken the glass mechanically. There are multiple sources of bubbles in conventional glass manufacturing, including air trapping, decomposition of the individual components that make up glass, galvanic oxidation reduction reactions, and “reboil”, which is the precipitation of super saturated glass when heated to a high temperature [22]. In our previous study of bubble formation in soda lime glass, three different forms of bubbles 998were identified in printed soda lime glass; periodic bubbles, sporadic bubbles and foam layers [20].

The foam layers were likely a result of reboil due to over heating the glass with the laser, while the periodic and sporadic bubbles could be attributed to bubble entrapment at the lower interface of the printed track and bubble generation around defects or contamination in the filament, respectively.In this paper, bubble formation in additive manufactured borosilicate glass is investigated. Two types of bubbles are identified for this glass system; filament defect/interface and reboil/volumetric defects. The physical phenomena associated with the two classes of bubbles are studied empirically. This helps understand the effects of the process parameters and leads to bubble-free borosilicate glass pieces. Finally, residual thermal stresses generated during the printing process are observed using a polariscope.

Finally, residual thermal stresses generated during the printing process are observed using a polariscope. These are removed using a post deposition annealing step without modifying the printed topography.Experimental SetupThe experimental setup for glass filament fed additive manufacturing process is shown in Fig. 1. A CO2 laser beam is incident on the top surface of the printed part. A glass filament is continuously fed into the intersection of the workpiece and the laser beam. The laser heats this interface to create a molten region above the glass transition temperature. The wavelength of the CO2 laser is 10.6 μm, which directly couples to phonon modes in the glass. This leads to a very short optical penetration depth so that the laser is effectively a surface flux. The movement horizontal movement of the stage is realized by x and y stages (Aerotech ANT130- 160XY), and a z stage (Aerotech ATS100-150) is used to raise and lower the platform.

The movement horizontal movement of the stage is realized by x and y stages (Aerotech ANT130- 160XY), and a z stage (Aerotech ATS100-150) is used to raise and lower the platform. The filament is fed by a homebuilt feeder. 2 mm diameter borosilicate glass rods are used in this paper (McMaster Carr).An unheated ceramic refractory block is used as the substrate. For small pieces of borosilicate glass printing, preheating of substrate to avoid thermal shock is not necessary because the CTE of borosilicate glass (~3×10-6 K-1) is only one third of that of soda lime glass. For larger printing larger pieces, the substrate may be needed to be heated to release the thermal stress. The laser is defocused to a Full Width Half Maximum (FWHM) diameter of 2 mm at build-plane filament intersection. A thermopile power meter (Ophir 10A–V1) is used to measure the reflected laser beam from the beam splitter, which is 1% of the laser energy delivered to the platform.

A thermopile power meter (Ophir 10A–V1) is used to measure the reflected laser beam from the beam splitter, which is 1% of the laser energy delivered to the platform. .1ShutterLaser Power, PFeed Rate, fCO2 LaserPower MeterPrinted GlassRefractory BlockScan Speed, vx-y-z Stage‘ Fig. 1. Illustrate of filament fed glass AM setupResults and DiscussionFigure 2 shows a dark field image of a printed piece of borosilicate glass. There are several clusters of bubbles visible in the specimen. Close observation reveals that the bubble clusters fall into two classes: (1) tightly spaced bubbles at the interfaces between tracks and (2) volumetrically distributed bubbles.

Close observation reveals that the bubble clusters fall into two classes: (1) tightly spaced bubbles at the interfaces between tracks and (2) volumetrically distributed bubbles. Microscope images of the interface concentrated bubbles show that their typical diameter is on the order of 10 µm while the volumetric bubbles are less densely spaced and have bubble diameters ranging from 50 to 100 µm.999(a)(b) Volumetrically distributed foam layer1 mm(c) Bubbles concentrated at interface1 mm5 mmFig. 2. An example of bubbles in a polished section of printed borosilicate glass (a) dark-field image of thewhole wall (b) foam layer of volumetrically distributed bubbles (c) bubbles concentrated at interface.Close inspection of the filaments revealed that many of them have scratches as received, prior to printing. These scratches are a possible source of bubble formation. To demonstrate this phenomenon, a diamond scribe was used to scratch a filament.

These scratches are a possible source of bubble formation. To demonstrate this phenomenon, a diamond scribe was used to scratch a filament. The filament was laid on the x-y-z stage as opposed to feeding the filament into the molten region. The laser was scanned along the along the filament with a power of 25 W and a scan speed of 0.5 mm/s. Figure 3 shows dark field images of the results. After laser irradiation, the scratch disappears and is replaced by clusters of small bubbles with diameters on the range of 10 µm. This supports the hypothesis that surface defects on the filaments are the origin of the tightly clustered bubbles in Fig. 2(c) which appear at the interface between tracks and have a similar size distribution.(a)5 mmScratch1 mm(b)BubblesFig. 3. Dark-field images of scratched filament (a) before and (b) after laser scanningSelecting filaments without scratches avoids the type of bubbles in Fig. 3 and Fig 2(c).

3. Dark-field images of scratched filament (a) before and (b) after laser scanningSelecting filaments without scratches avoids the type of bubbles in Fig. 3 and Fig 2(c). Single track wide walls were printed using a constant scan speed and feed rate, v=f=0.5 mm/s. After each row, the z-stage was lowered 1 mm. This was based off measurements of single track heights for the 2 mm diameter borosilicate filaments.Six walls were printed Single walls could be printed from 25 to 50 W, while the filament could not be fully melted. All of the six walls are polished, and taken dark-field image to quantify the bubbles inside the samples. The bubble density of the walls is shown in Fig. 4. No bubbles were visible for low laser powers (≤30 W). At higher laser powers (≥35 W), foam layers could be observed. The average diameter of bubbles was larger at higher laser powers.

4. No bubbles were visible for low laser powers (≤30 W). At higher laser powers (≥35 W), foam layers could be observed. The average diameter of bubbles was larger at higher laser powers. Because of the removal of scratched filaments, no small bubbles at track interface were observed in any of the samples.The temperature of molten region increases with laser power. The observation that the bubble concentration scales with laser power suggests that rate of bubble formation scales with temperature. The experimental results also show that below a laser power threshold the glass does not break down, which is consistent with the reboil phenomenon in conventional glass production. Vapor rising from the molten region1000was observed for powers greater than 35 W which also supports a temperature driven chemical breakdown of the glass.350]3 -300m c [250y t i s n e de b b u Bl200150100500 253035404550Laser Power [W] Fig. 4.

4. Bubble density of the printed wallsResidual stresses are another important factor limiting the quality of printed glass. This issues is more significant for the borosilicate glass printed on an unheated platform than previous soda-lime work using a heated substrate. Beyond the risks of thermal shock, residual stresses are a significant issue for printed optics because they result in local index anisotropy [22] and reduced shatter resistance. While residual stresses are also a significant issue for many additive manufacturing processes, printing transparent glass offers a unique opportunity to study the phenomena. Stress induced birefringence can be detected using a polariscope [23].Figure 5(a) shows an image through a polariscope of a single track sample as printed (prior to polishing). The bright regions indicate stress concentration.

The bright regions indicate stress concentration. Future studies will explore the effects of path planning on the residual stress which is caused by uneven heating/cooling cycles during deposition. Annealing the sample can relieve trapped stresses. To demonstrate this, the sample in Fig. 5(a) was soaked in a 568°C in a muffle furnace for two hours before cooling to the strain point at 510°C over two hours and finally cooling to room temperature over the course of an hour. A polariscopic image of the sample after annealing is shown in Fig. 5(b). The stress- concentration is significantly reduced. It is important to see, that the morphology of the sample was not significantly affected by the annealing process.(a)(b)5 mmFig. 5. Polariscope images of a printed borosilicate glass wall (a) before annealing (b) after annealingThe furnace annealed wall Fig 5(b) was then polished on both sides. The dark-field image of the polishedwall is shown in Fig. 6.

The dark-field image of the polishedwall is shown in Fig. 6. No bubbles are observed in the sample and the birefringence is clearly reduced.5 mmFig. 6. Polariscope image of polished bubble-free glass wall after polishing1001Summary and ConclusionThis paper explored additive manufacturing of borosilicate glass. Two different bubble regimes are identified in the printed borosilicate glass: small bubbles at track interface and foam layers. The formation small bubbles at track interface is traced to surface defects on the filament. The foam layers are traced to the reboil of glass at high temperatures. Bubble free glass can be produced by avoiding scratches on the filaments and minimizing the high temperature in the molten region by printing slower with lower laser power or using smaller diameter filaments. Significant residual stresses were observed in the printed pieces.

Significant residual stresses were observed in the printed pieces. These stresses were able to eliminated with a post annealing process without changing the overall morphology of the printed pieces. The elimination of bubbles and residual stress induced birefringence is key to printing glass for optical applications or low CTE structural elements.AcknowledgmentThis work was supported by the National Science Foundation (CMMI-1538464).References[1] Gibson, I., Rosen, D. W., and Stucker, B., 2010, Additive Manufacturing Technologies, Springer. [2] Brockmeyer, E., Poupyrev, I., and Hudson, S., 2013, "PAPILLON: designing curved display surfaces withprinted optics," Proc. ACM Symposium, St. Andrews, Scotland, pp. 457-462.[3] Willis, K., Brockmeyer, E., Hudson, S., and Poupyrev, I., 2012, "Printed optics: 3D Printing of EmbeddedOptical Elements for Interactive Devices," Proc. ACM Symposium, Cambridge, MA, USA, pp.

214-219.[13] Fateri, M., and Gebhardt, A., 2015, "Selective Laser Melting of Soda-Lime Glass Powder," InternationalJournal of Applied Ceramic Technology, 12(1), pp. 53-61.[14] J. Luo, H. Pan, and E. C. Kinzel, 2014, “Additive Manufacturing of Glass,” Journal of ManufacturingScience and Engineering, 136(6), pp. 061024-061024.[15] Fu, Q., Saiz, E., and Tomsia, A. P., 2011, "Bioinspired Strong and Highly Porous Glass Scaffolds,"Advanced Functional Materials, 21(6), pp. 1058-1063.[16] Klein, J., Stern, M., Franchin, G., Kayser, M., Inamura, C., Dave, S., Weaver, J. C., Houk, P., Colombo, P., and Yang, M., 2015, "Additive Manufacturing of Optically Transparent Glass," 3D Printing and Additive Manufacturing, 2(3), pp. 92-105.1002[17] Luo, J., Gilbert, L., Qu, C., Wilson, J., Bristow, D., Landers, R., and Kinzel, E., 2015, "Wire-Fed Additive Manufacturing of Transparent Glass Parts," Proc.

viewed PaperBuild Plate Design for Extrusion-Based Additive ManufacturingAlex Roschli, Brian K. Post, Celeste Atkins, Adam G. Stevens, Phillip Chesser, Kristin ZaloudekOak Ridge National Laboratory, Manufacturing Science Division, 2350 Cherahala Blvd, Knoxville, TN 37932AbstractA fundamental part of machine design for large format polymer extrusion-based additive manufacturing (AM) systems is the substrate where the object is to be constructed, often referred to as the build plate. A good build plate is imperative for a successful build as it is used for supporting and positioning the part during the entire construction. For planar 3D printing, this build plate needs to be flat and in-plane with the X/Y motion of the 3D printer. Additional functionality can include heaters for preventing delamination or warping, vacuum to help prevent warping, removable build surfaces for quick part removal, and mounting features for helping position a part for subtractive operations.

This paper reviews existing build plate designs and discusses the design considerations and materials for build plate fabrication.Keywords: build plate, build surface, extrusion, 3D printer designIntroductionExtrusion-based additive manufacturing works by depositing material in layers onto a build surface [1-3]. The build surface can be a sheet that sits atop a build plate, or the surface and plate can be the same object in which case the surface would reference the top surface of the plate. Throughout the context of this paper, build surface shall refer to the area where printing occurs and will not be used to reference the entire assembly. The term build plate will be used to refer to a plate supporting the build surface or in reference to the entire assembly in contexts where a separate build surface is not needed.The build plate and build surface are very important for successful part construction, and without them, the build will likely fail.

Properly designed and implemented build plates and surfaces have many characteristics including: good bonding/adherence to the first printed layer, near perfect flatness in-plane with the X/Y motion, and the ability to support the load of the entire printed object.The build surface is needed to bond and adhere the first layer during the deposition process to lock the build in place [4]. It is imperative that the first layer be printed and captured within the same space for the entire duration of the print. If the material being extruded does not stick on the1079first layer, it could drag around the surface and create build-up on the nozzle. This failure to stick on the first layer almost certainly means a failure of the entire build because subsequent layers will not properly adhere and take the correct shape to support future layers. The build surface needs to not only adhere to the first layer during deposition but needs to remain adhered for the entire duration of the print.

The build surface needs to not only adhere to the first layer during deposition but needs to remain adhered for the entire duration of the print. If the first layer begins to warp or peel away from the build surface, the entire part can become distorted and lose dimensional accuracy. If the entire first layer separates from the build surface, the object may move freely about the build volume and future layers will not adhere correctly causing a failed build. However, at the conclusion of the print, it is desirable for the build surface to easily separate from the part and be reusable for future builds. For thermal- based processes, such as thermoplastic extrusion, this would mean having the build surface adhere to hot plastic and separate easily when cooled.For traditional 3-axis 3D printing, wherein layers are built in the X/Y plane, then the Z height is advanced to create a new layer, the build surface must be flat and in-plane with the X/Y motion of the system.

A non-flat build surface, particularly referring to flatness with respect to the printer’s X/Y motion, is very apparent on the first layer because certain areas of the print will stick while low areas may be extruded above the sheet and high areas can cause the nozzle to collide with the surface. Perfect flatness is not a requirement, due to high cost and complexity with maintaining a perfectly flat surface. However, flatness within 25% of the first layer height is desirable. This 25% or less variation in flatness is desirable because small amounts of variation in the surface can be absorbed by the material flow during the extrusion process. After the first layer, if the build surface does not move, the top of the part will be in-plane with the X/Y motion and subsequent layers will be flat with respect to one another.A proper build plate must also be able to support the load of the part being constructed.

With each successive layer that is deposited, the total weight, and thus the total load on the build plate, grows. Modern large format printers can produce objects the size of an automobile with total masses exceeding 1000kg. If the build plate deflects under this load, the top surface of the print may stop being flat and cause adhesion issues for future layers. On larger printers, it is possible for the printed object to become stiffer than the build plate meaning that the warping of the build plate is less obvious at the top layer of the printed object; however, this may create the issue of the build surface and part separating from one another which allows the part to move.The following sections will review existing designs that are currently available in the marketplace.

Then, an analysis will be provided for specific design strategies and aspects of a good build plate.Review of Existing DesignsCommercially available large and small format extrusion systems offer many different machine configurations and build plate designs with varying functionality including heated vacuum sheet tables, heated and vacuum taped sheet tables, wood board with glued pellets, heated glass sheets, and a variety of off-the shelf build sheet products such as painters’ tape and adhesive sheets. While all these designs have their merits, some may not scale well to large format processes. Each design will be reviewed, and the following section will then take a deeper look at1080some of the design strategies at play to understand different means of successfully designing a build plate.Heated vacuum sheet tables use a build sheet, typically made from a material that will bond well with the printed material [5], as a build surface for securing the part during printing.

This type of setup is commonly seen on polymer filament-based extrusion processes where residual stress and warping create small forces that will not overcome the strength of a pneumatic vacuum that is being used to hold the sheet in place. The heat is used to keep the sheet and first few layers of the part warm, ideally at the glass transition temperature (the temperature at which the polymer begins to soften) of the polymer, so that the printed part stays relaxed and flat during the entire duration of the print. A vacuum, typically distributed through a grid in the build plate or a window screen placed on top of the build plate, is used to hold the sheet in place during the print. It is sealed on the outside edges by one large continuous O-ring. The advantage of this approach is the ease of setup. A build sheet can very quickly be placed inside the machine and vacuumed down in place. When the print is complete, the vacuum is released, and the print is removed.

A build sheet can very quickly be placed inside the machine and vacuumed down in place. When the print is complete, the vacuum is released, and the print is removed. A new sheet can quickly be added, and printing can immediately resume. Depending on the sheet and printing materials, sheets can often be reused after a part has been removed. The disadvantages are that prints with a lot of warping and curling can cause the build sheet to lose vacuum during the print and no longer be constrained, the solution does not scale well, and a constant vacuum force is needed for the duration of the build. A large sheet requires a lot of vacuum force to hold it down and is very likely to lose seal at one edge.Vacuum tables with taped sheets differ from vacuum sheet tables in that they use tape around the build sheet edges to maintain a vacuum seal. This type of setup is commonly used for large format polymer processes where the large prints create high residual stresses and warping.

This type of setup is commonly used for large format polymer processes where the large prints create high residual stresses and warping. The tape for these setups, usually made from a vinyl-based material, is designed to stretch if the build sheet flexes so that vacuum is still maintained. Often during printing, the residual stress formation will cause the part, and sheet it is attached to, to start warping. Without tape, the sheet may pull away from the vacuum table and thus lose vacuum, leaving it free to continue warping. The tape helps maintain that vacuum seal and allows the vacuum to continue pulling down on the part to help keep it flat. This design requires a vacuum distribution grid to ensure that the vacuum force makes it all the way to the edges, or a vacuum spreader, such as a window screen, to be used so that vacuum force makes it way to the tape. The advantages of this setup are the scalability and reusability.

The advantages of this setup are the scalability and reusability. The addition of tape allows for large sheets (upwards of 2 meters on a side) to be held in place on long prints without losing vacuum. Tape also allows for thicker build sheets to be used which are more likely to withstand the print removal process and thus can be reused multiple times. The disadvantages are the labor and material cost involved with applying tape all the way around the sheet as well as the need for a constant vacuum force during printing. Figure 2 shows a Cincinnati Big Area Additive Manufacturing (BAAM) system with an acrylonitrile butadiene styrene (ABS) plastic build sheet taped on the edges and held in place with a vacuum.

Figure 2 shows a Cincinnati Big Area Additive Manufacturing (BAAM) system with an acrylonitrile butadiene styrene (ABS) plastic build sheet taped on the edges and held in place with a vacuum. This system has a single vacuum port and uses a window screen to help diffuse the vacuum.1081Figure 1: Cincinnati BAAM system with taped build sheets on a vacuum table.Wood, or other solid surface, with glued pellets is a relatively new solution used by large format polymer processes. The typical setup involves a sheet of medium-density fiberboard (MDF) coated with a mixture of wood glue and polymer pellets. The glue allows the pellets to adhere to the surface for the duration of the print. The pellets are the same polymer that is used for the printing process so that a strong bond is achieved between the pellets and print. The advantages of this setup are that the setup works for large parts and parts can be mounted to the board to prevent warping.

The advantages of this setup are that the setup works for large parts and parts can be mounted to the board to prevent warping. Build sheets, like those used for the two previous designs, only come in certain sizes meaning that multiple vacuum tables are needed to hold sheets for large prints. With this setup, an unlimited number of boards can be coated with pellets and set alongside each other. With the build surface being wood based, it’s easy to use woodscrews to screw the print into the board during printing to help combat any warping or delamination. The disadvantage of this setup is the cost and time to setup and use a new board before each print. Figure 2 shows the patented bead board approach from Thermwood [6].Figure 2: Thermwood’s version of a wood board with glued pellets, called “bead board” [6]1082Heated glass sheets are a simple and low-cost solution ideal for small polymer filament- based extrusion processes.

A solid glass sheet, usually made from borosilicate glass because of its improved resistance to cracking at high temperatures, is usually set atop a heated metal build plate and held in place with binder clips or tabs. The polymer can be printed directly on the clean glass when hot, and then easily removed when cool. Not all polymers adhere to the glass on their own, but adhesion can be improved with products such as hairspray, glue sticks, and polymer dissolved in acetone. Commercially adhesion promotors such as Stick Stick and Magigoo are also available. The advantages of this setup are the low cost and simplicity of setting it up. The main disadvantage is the difficulty with getting the print to properly adhere to the glass sheet which is why the market for adhesion promotors exists. However, regular use of an adhesion promoter can lead to a buildup of a glue-like substance on the printer that has to be cleaned away. Another disadvantage is scalability.

However, regular use of an adhesion promoter can lead to a buildup of a glue-like substance on the printer that has to be cleaned away. Another disadvantage is scalability. Glass sheets work well for small parts, but larger parts become too heavy with too much residual stress such that a glass sheet is ineffective. Figure 3 shows a Makergear M2 printer with a glass build sheet held in place by binder clips.Figure 3: A Makergear M2 with glass build surface.A final design is an adhesive sheet or tape on a solid surface that has no vacuum. This setup can be heated, or not heated, and is commonly used for small polymer filament-based extrusion processes. Painters’ tape is a cheap and easy to implement solution that can be acquired at most local stores. Strips of the tape are applied directly to the flat surface and printing occurs directly onto the tape. This works well with low temperature materials such as PLA (polylactic acid).

Strips of the tape are applied directly to the flat surface and printing occurs directly onto the tape. This works well with low temperature materials such as PLA (polylactic acid). For higher temperature materials, and those with a high coefficient of thermal expansion (CTE), which makes warping more likely, special 3D printing sheets can be used. These sheets are typically adhesive backed and can be attached directly to a build plate so that they are fixed in place. Some systems use a magnetically attached flex-steel build plate with an adhesive sheet on top allowing the user to quickly remove and flex the sheet to dislodge the printed part (Figure 4). The coatings1083on these sheets are often proprietary, but some simply use a thin layer of polyetherimide (PEI) polymer because of its adhesive properties. Like the heated glass sheet setup, the advantages are low cost and simplicity of setup. The disadvantage is scalability.

Like the heated glass sheet setup, the advantages are low cost and simplicity of setup. The disadvantage is scalability. Large adhesive sheets aren’t readily available and may not provide enough hold-down force for materials with a high coefficient of thermal expansion. Flex-steel based implementations don’t scale because of the difficulty in removing and flexing such a large heavy sheet.Figure 4: Prusa Mini Build Plate (left) and Prusa Mini Build Sheet (right). In this setup, the build sheet is made from flex steel and coated with a PEI sticker. The build sheet magnetically attaches to the build plate.Design StrategiesHeated Versus Unheated Build SurfacesA first major design decision for creating build plates is whether to heat the build surface. A heated build surface is beneficial because the heat helps keep the first layers of the print at an elevated temperature, which prevents the polymer from cooling and contracting which often leads to warping and delamination.

Many researchers have proven the benefits of a heated substrate to improve dimensional accuracy, material strength, and print quality [7-10].The implementation of a heated build surface is similar on both large-format and small- format thermoplastic extrusion AM systems. The most common setup uses single-point temperature feedback for build surface heating, with the build area divided into one or more single- point feedback zones depending on the size of each heater element relative to the build area. Electrical resistive heaters are most common, AC or DC depending on the system scale, with AC systems favored at higher powers due to infrastructure considerations.

Electrical resistive heaters are most common, AC or DC depending on the system scale, with AC systems favored at higher powers due to infrastructure considerations. Common heater designs include silicone-glass composite resistive heaters [11], which are adhered to the underside of the build substrate; and circuit board heaters which incorporate resistive traces directly into a printed circuit board [12,13], which are alternatively adhered to or mechanically clamped against the build surface. It is common for heater underside to be insulated to improve efficiency by reducing heat loss to the environment. In the simplest case the insulation consists of the PCB board itself, larger installations use glass fiber matting or similar insulation.1084While many researchers have documented the value and improvements from a heated bed, the unheated bed can still have its merits such as reduced cost.

Heating elements for a build surface add cost to the overall system, not just in the cost of the heaters themselves, but in the cost of all the other additional components needed such as heater controllers, wiring, and temperature sensors. For polymer extrusion printers, the heated bed uses as much, or more, power than the rest of the system components combined. There is also the added complexity of integrating the heaters and properly insulating and isolating the heaters from the rest of the system. The heater placement has to be properly designed so that the heat is directed into the build surface and not into the other components of the system, such as motors and drives which are sensitive to heat. If the heaters aren’t properly insulated, they will use more electricity to maintain temperature, which will cause the build environment to heat up. A heated chamber is desirable [14], but only if the other system components can handle operation and high ambient temperatures.

A heated chamber is desirable [14], but only if the other system components can handle operation and high ambient temperatures. With all these considerations in mind, it can be beneficial to forego the heated bed whenever the process allows. Low temperature materials, particularly those with a low CTE, can be printed on heated surfaces when the other process parameters are optimized [15].Design of Vacuum Grooves and Hold-Down FeaturesAs previously mentioned, there are many styles of build plates that use vacuum fixturing, or the process of using a distributed vacuum to hold a build surface or work piece in place during manufacturing. This vacuum arrangement requires a distribution and edge sealing method.Often the build surface is made of a relatively thin, flexible material such as the ABS build sheets commonly used on BAAM.

This vacuum arrangement requires a distribution and edge sealing method.Often the build surface is made of a relatively thin, flexible material such as the ABS build sheets commonly used on BAAM. When only a single vacuum port is used for a flexible sheet, it's possible for the sheet to deform such that a vacuum seal is created around the port hole which prevents vacuum from dispersing to the rest of the sheet. Increasing the number of holes in the plate will increase the area that is being sealed, but the spacing would need to be very close together requiring a large number of holes and additional fittings for the air lines. This is not a very economical solution. With a method of distributing the vacuum, depending on the size, a build plate can use just one vacuum port and still apply adequate suction.A window screen material or mesh can work well to distribute the vacuum force.

It creates tiny gaps between the build surface and the underlying plate to allow the suction to be distributed all the way to the edges. The same goal can be accomplished with a shallow grid, as shown in Figure 5, or pattern machined into the surface of the supporting plate. While it depends on the flexibility of the build surface material, a narrow groove, likely 1x – 3x the thickness of the build surface, can be spaced at approximately 10cm on center in a grid pattern for suctioning down the entire build surface.1085Figure 5: Vacuum distribution system using grid and sealWith either of these methods, the edge needs to be sealed in some way. It was previously discussed that some systems use O-rings to seal the outer edge of the build surface to the plate, but this does not work well for large format systems. Taping the edge of the sheet to the plate works consistently, but the plate must be designed appropriately.

Taping the edge of the sheet to the plate works consistently, but the plate must be designed appropriately. To use tape for edge sealing, the grid should not extend all the way to the edge, and there should be space between the edge of the plate and the outer edge of the grid that is smooth and uninterrupted for the tape to stick to. Three centimeters is adequate in most applications. Tall soft closed cell foam seals and flange mounted rubber seals have also been investigated with favorable results. To use these types of seals, a properly sized machined grove to receive the seal is required. This groove should be a closed loop that does not connect to the outside of the plate.Vacuum fixturing is a well-known method of part fixturing for traditional manufacturing. These systems often see small amounts of side loading and distortion without failure, and the same methodology can be applied to hybrid manufacturing systems if side loading and distortion are minimized.

For instances when that is not the case, some additional clamping fixture is necessary. T-slots, shown in Figure 6, are commonly used to fixture parts for machining with toe clamps. These can be added to the edges of the build plate to facilitate the use of bar clamps for holding the part in place.1086Figure 6: T-slot in build plate for clamping featuresBuild Plate MaterialThe material used to make the build plate is another important characteristic that should be considered during design. Desktop-scale systems produce parts on the order of 3kg or less. This reduces the load bearing requirement for the build plate, which is commonly sheet metal [16] or the fiberglass printed circuit board heater, either alone or in conjunction with a rigid build surface such as plate glass [17]. The resulting low thermal mass of the bed assembly may result in temperature nonuniformity across the bed due to natural convection and result in part warping.

The resulting low thermal mass of the bed assembly may result in temperature nonuniformity across the bed due to natural convection and result in part warping. Solutions incorporating gradient resistive elements throughout the PCB have been developed, to provide additional heat output at the build plate boundary condition [18].To support the greater mass of parts printed on large-format systems (on the order of 10- 1000 kg), build plates are commonly manufactured from metal plate, typically aluminum for a compromise between stiffness, thermal conductivity, and cost. Similar to the desktop-scale printers, electrical heaters can be fixtured to the underside via adhesive or mechanical clamping. Build plate thickness is dictated by the requirement to provide sufficient stiffness (a function of plate thickness and spacing of underlying support points) while maintaining an approximately uniform temperature distribution within the body of the plate.

This is required because the temperature feedback is typically located on the plate underside, immediately adjacent to the heaters, and the print occurs on the topside. For such a lumped-capacitance model to be valid, the Biot number must be kept less than approximately 0.1. This ensures that the temperature feedback provided by the under-bed sensor accurately reflects build surface temperature. The Biot , where h is the convective heat transfer coefficient, k is the body number is defined as thermal conductivity, and L is the characteristic length of the body, which, in the case of a large plate held at constant temperature on one side, is equal to the plate thickness. The upper limit on , assuming sufficient heater power to build plate thickness at steady-state is therefore hold the heated side of the plate at a constant temperature. Bergman et al.

The upper limit on , assuming sufficient heater power to build plate thickness at steady-state is therefore hold the heated side of the plate at a constant temperature. Bergman et al. [19] provides a method to estimate the convective heat transfer coefficient of a horizontal plate immersed in air with ) [20], the maximum thickness is 3m, whereas natural convection. For aluminum (ℎ 𝑘𝑘 𝐿𝐿𝐵𝐵𝐵𝐵 =𝑘𝑘 ℎ𝐿𝐿 = 0.1𝑊𝑊) [21] it is 0.2m. In practice, achieving the lumped parameter condition is for steel ( 𝑚𝑚−𝐾𝐾 trivial, and build plate thickness is commonly on the order of 2.5cm on large format systems. This𝑘𝑘𝐴𝐴𝐴𝐴 = 239𝑊𝑊𝑘𝑘𝑆𝑆 = 15𝑚𝑚−𝐾𝐾1087design requirement to satisfy the lumped-parameter condition also enables accurate single-point temperature feedback of first layer process temperature if heat is to be drawn out of the part at steady state, e.g.

in the case of an exothermic curing reaction in the printed bead.While metal is the most common build plate material, wood has been implemented on some systems. Wood is a great material because it is readily available, cheap, and does not conduct heat well which prevents it from drawing heat from the part. However, wood does not provide enough stiffness for large and heavy prints, and the poor thermal conductivity also prevents wood from conducting heat from the heating elements to the part. Systems can overcome the stiffness issues by designing a fixed table, making it easier to support on the underside. For lower temperature materials, such as PLA, a heated plate is not required so long as the plate doesn’t pull heat away from the part, making a wood plate a good fit for PLA.ConclusionA properly designed build plate and build surface are paramount for conducting a successful 3D print.

The build surface must adhere to the first layer, maintain flatness within the X/Y plane, and be able to support the load of the full object during the entire printing process. Many different solutions exist in the marketplace today, but there is no one size fits all solution because different processes and machines have different requirements. Certain solutions, such as glass plates and flex steel sheets, do not scale as well as others such as polymer build sheets and vacuum tables. Polymer build sheets typically require heat to promote adhesion which requires a solid support surface that can effectively transfer the heat to the sheet. An effective vacuum table needs to have proper vacuum distribution and room for tape adhesion. Additional features such as a vacuum grooves and T-slot grooves can also be implemented.

An effective vacuum table needs to have proper vacuum distribution and room for tape adhesion. Additional features such as a vacuum grooves and T-slot grooves can also be implemented. Further research is needed to improve existing solutions and develop new solutions to keep up with the growing size of 3D printers.AcknowledgmentsThis material is based upon work supported by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Office of Advanced Manufacturing, under contract number DE-AC05-00OR22725.References 1. Shaqour, Bahaa, et al. "Gaining a better understanding of the extrusion process in fused filament fabrication 3D printing: A review." The International Journal of Advanced Manufacturing Technology 114.5 (2021): 1279-1291.2. Kumar, Sudhir, et al. "Fused filament fabrication: a comprehensive review." Journal of Thermoplastic Composite Materials (2020): 0892705720970629.3. Soni, Shivam, and Mohammad Taufik.

ild Style Decision Support for StereolithographyJoel E. McClurkin and David W. Rosen Rapid Prototyping and Manufacturing Institute Georgia Institute of Technology Atlanta, GA 30332-0405(404) 894-9668http://rpmi.marc.gatech.eduABSTRACTWhen building parts in a stereolithography apparatus (SLA), the user is faced with many decisions regarding how the part will be built. The quality of the build can be controlled by the user via changing one of several build style variables, including part orientation, cross sectional layer thickness, and laser hatch density. A user will probably have preferences for the part build (Le., accuracy or speed), but may not understand how to vary the build style variables to produce the desired results.

A user will probably have preferences for the part build (Le., accuracy or speed), but may not understand how to vary the build style variables to produce the desired results. A method based on response surface methodology and multi-objective decision support is described in this paper for relating build goals to three build style variables, and the use of these relationships in providing decision support for building a part on a SLA. The method is applied to the build style of a circuit breaker handle.INTRODUCTION Stereolithography (SLA) machines have dozens of processing variables that can be controlled by experienced operators to meet exacting build requirements of accuracy, surface fmish, and build time. Experienced operators know qualitatively how process variables are related to build goals and can quantify some of their knowledge.

Experienced operators know qualitatively how process variables are related to build goals and can quantify some of their knowledge. Although SLA machine operation is predictable, the quality of resulting parts is not always obvious, particularly when trade-offs must be made among accuracy, surface finish, and build time.1.A decision support method is presented in this paper to aid SLA operators in selecting appropriate values of process variables in order to achieve build goals as closely as possible. The goals can have importances specified in order to reflect preferences of the operator or part designer. The method is adapted from multiobjective optimization and utilizes response surface modeling to quantitatively relate variables to goals. Since our study was limited, we looked at only three process variables: part orientation, layer thickness, and hatch spacing, and three goals: accuracy, surface finish, and build time.

Since our study was limited, we looked at only three process variables: part orientation, layer thickness, and hatch spacing, and three goals: accuracy, surface finish, and build time. Designer/operator preferences among the goals are also modeled using two different methods: weightings and priorities. Our objective is to render decision support by handling trade-offs among conflicting objectives quantitatively. Our method is demonstrated on a part with non-trivial geometries, a circuit breaker handle. Results reflect our intuition.2. PREVIOUS RESEARCH 2. 1 SLA Build Style OptimizationThere have been several approaches to optimizing the build performance of SLA machines. These approaches vary from finding an optimal build layout to reduce build time to adaptively slicing a part to improve surface finish and reduce build time. One approach for optimizing build layout has been to use an expert system for determining preferred build direction as shown by Frank and Fadel [1].

One approach for optimizing build layout has been to use an expert system for determining preferred build direction as shown by Frank and Fadel [1]. Expert systems are heuristic in nature because most of the rule systems and knowledge used to create them are based on the experience of an expert. For example, an expert knows that thin layered parts build slower on the SLA, but are more accurate. This expert system uses information on geometric feature orientations to determine the preferred build style.Another method for automating build layout addresses the problem of laying out a build with multiple parts to minimize build time. A problem arises when the parts need to be packed as closely as possible to maximize the utilization of space in the build vat and also reduce build time.627A genetic algorithm was developed by Wodziak et al. [2] to automatically place multiple parts to reduce build time.

[2] to automatically place multiple parts to reduce build time. The parts may be translated or rotated +/- 90 degrees about the z-axis to aid in part packing. The creators of the genetic algorithms for part packing reported it to be a successful method for optimizing multiple part layout to reduce build time.Build style optimization research has also been done with regards to build time and surface finish of curved surfaces. A technique, called adaptive slicing, involves slicing the ·part thinner in areas of high curvature and thicker throughout flat areas [3]. The main idea is to estimate the surface finish for a particular curved· surface by detennining how much stair stepping is present. If the amount of stair First, the part is sliced using the maximum allowable layer thickness. stepping creates a surface that is rougher than a pre-determined limit, the part is sliced thinner in that region.

stepping creates a surface that is rougher than a pre-determined limit, the part is sliced thinner in that region. This work was done in the context of a fused deposition modeler (FDM).A build style issue that is not dealt with in this paper, but is very important, is support structure generation. A method has been developed by Hur and Lee [4] to automatically generate support structures in SLA build styles. Not only does their method generate support structures, but it determines the best build orientation to minimize the amount of support structures required. Hur and Lee determine which facets of the STL file need to be supported based on the following criteria: orientation of the facet, area of the facet, part instability, and base support.The build optimization research discussed in the previous paragraphs is based primarily on single objective optimization. For instance, the genetic algorithms for part packing utilized the single objective of reducing build time.

For instance, the genetic algorithms for part packing utilized the single objective of reducing build time. The research in this paper deals with multiobjective decision support, where several objectives such as build time and accuracy are considered. A companion study was performed, based on selection among alternatives rather than multiobjective optimization, and was reported in [5].2.2 The Compromise DSPThe method of rendering build style decision support described in this paper is based on solving a compromise Decision Support Problem (DSP) [6]. The compromise DSP is a type of multiobjective optimization method and is used to model decisions that consist of decision variables, constraints, and multiple goals. The goals are often in conflict with one another (i.e., accuracy and build time). The compromise DSP is an extension of Goal Programming and Sequential Linear Programming [7].

The goals are often in conflict with one another (i.e., accuracy and build time). The compromise DSP is an extension of Goal Programming and Sequential Linear Programming [7]. Solution of compromise DSPs is typically performed using the Adaptive Linear Programming (ALP) algorithm. The structure of the compromise DSP is shown below.Given: Find: Satisfy: Minimize:A feasible alternative, assumptions, parameter values. Values of design and deviation variables. System Constraints, System Goals, and Bounds on variables. Deviation Function that measures distance between goal targets and design point.To use the compromise DSP, a part is presented in a "default" build style to serve the purpose of the existing alternative to be improved. This build style is then improved by changing the build style variables. The goals are build speed, part surface finish, and part accuracy. More accurate builds usually take longer to build.

This build style is then improved by changing the build style variables. The goals are build speed, part surface finish, and part accuracy. More accurate builds usually take longer to build. The alternative (in this case a build style) is improved by finding a combination of system variables such that all the system constraints are satisfied while the deviation function is minimized. The deviation function is solely a function of the deviation variables, which in tum are a measure of how well the system goals are met.3. BUILD STYLE DECISION SUPPORT METHODNow that some of the background pertaining to build style decision support has been presented, the methods for solving the build style decision support problem are presented. This includes a brief discussion of the important build style variables and goals and how they are applied in the formulation of a compromise Decision Support Problem.6283.

This includes a brief discussion of the important build style variables and goals and how they are applied in the formulation of a compromise Decision Support Problem.6283. 1 Build Style Variables, Parameters, and GoalsThe build style variables, parameters, and goals are those quantities which drive the solution of the compromise nsp. Build style variables are those quantities which describe the attributes of the build style. These variables are directly affected by user input. Throughout the remainder of this paper, the build style variables will also be referred to as system .variables. The system variables chosen for this project are the most common variables the user can change when setting up a build style. They are: Build Orientation (Surface Angle and Z-Height), Layer Thickness, and Hatch Spacing. Build Orientation The build orientation variable is important because it can have an effect on every measure of build performance.

Build Orientation The build orientation variable is important because it can have an effect on every measure of build performance. The important aspects of the build orientation are the z-height of the part, and the various angles the surfaces of the part create with respect to the x-y plane. The part z-height can be controlled by the user by changing the orientation of the part. The bounds for this variable are 0.0025" and 10", which are limited by the smallest layer thickness and the size of the vat, respectively. However, the bounds of this variable can be narrowed significantly by dimensional information contained in the STL file that represents the part. Z-height can affect build speed dramatically.

However, the bounds of this variable can be narrowed significantly by dimensional information contained in the STL file that represents the part. Z-height can affect build speed dramatically. Builds with larger z-heights take longer to build because there are more build layers to build, causing more elevator movement and delay times.Z-height also affects accuracy because the SLA is generally not as accurate in the z-direction due to overcure, print through, and build quantization errors [8]. Orientation also affects the locations of support structures which affect accuracy and surface finish. However, support structure generation and location is another issue too big to be dealt with in this paper. Technically build orientation can be considered a continuous variable due to the infInite number of orientations in which a part can be placed.

Technically build orientation can be considered a continuous variable due to the infInite number of orientations in which a part can be placed. Layer Thickness When the STL fIle is sliced by the Maestro™ [9] processing software, it is sliced into many cross-sectional layers of finite thickness. The layer thickness variable refers to the thickness of these layers. The user has the ability to specify the thickness of the layers that represent the sliced part. The bounds for this variable are 0.0025" to 0.02". These bounds were determined from the Maestro™ build software version 1.8.Layer thickness has a great effect on surface finish and accuracy, and to some extent build time. The accuracy and surface finish of curved surfaces will be better when thinner layers are used. However, parts usually build faster using thicker layers because recoat time will be greatly reduced. The phenomenon of build quantization plays a major role in accuracy and surface finish.

However, parts usually build faster using thicker layers because recoat time will be greatly reduced. The phenomenon of build quantization plays a major role in accuracy and surface finish. If the top of a part falls on a half layer thickness, the SLA builds the extra layer which adds an extra half layer thickness to the top of the part. The scan speed of the laser, coupled with the pre It is for the reasons above it is often necessary to dip delay, affects down facing z-errors [8]. select the correct value for layer thickness to increase build performance. Hatch Spacing Hatch spacing is important because it affects build time, part accuracy, and surface finish. Hatch spacing refers to the density of the hatching pattern the SLA uses when "filling in" each cross section of the build. When the hatch vectors are closer together, the laser must scan more vectors per unit of cross sectional area, and the part will take longer to build.

When the hatch vectors are closer together, the laser must scan more vectors per unit of cross sectional area, and the part will take longer to build. The fewer vectors the laser has to trace, the quicker the build will be completed. However, increased hatch vector density results in more resin being cured during the build and less during the post-cure process. This reduces part shrinkage, therefore increasing part accuracy. Also, as the hatch spacing on the up-facing surfaces (called fill vectors) increases, the quality of the surface finish likely crash, which was decreases. experimentally observed. The bounds for h~tch spacing were 0.004" to 0.020". Build Style Parameters The build style parameters are quantities which describe properties inherent to the part being built. The build style parameters will also be referred to as system parameters throughout the rest of the paper.

The build style parameters will also be referred to as system parameters throughout the rest of the paper. The user cannot change these properties because theyIf the hatch spacing is too wide,the build will most629are determined from the part's STL file or other uncontrollable parameters such as resin properties or laser power.3.2 Modeling Of Build GoalsThe build goals considered for this paper are: (1) Low Build Time, (2) High Accuracy, (3) In order for the build style to be improved, it is necessary to be able to Smooth Surface Finish. predict how the measures of build perfonnance (the system goals) will change as the build style variables are altered. Response Surface Modeling is used to model the behavior of two goals, build time and accuracy, as a function of the build variables and parameters.

Response Surface Modeling is used to model the behavior of two goals, build time and accuracy, as a function of the build variables and parameters. An analytical equation is used to predict the behavior of the other goal, surface finish.A response surface model is a means for expressing the build goal (the response) as a function of several build style variables. This is done by running several experiments and fitting a polynomial equation, called the response surface, to the resulting data. In support of this research, a series of experiments were conducted to generate data for the build time and accuracy response surfaces [10]. Four experimental factors were selected. A face-centered, central composite experiment design was used to create quadratic response surfaces. Block-shaped and cylindrical parts were built in various sizes and orientations to correspond to the experimental factors and their settings, shown in Table 1.

Block-shaped and cylindrical parts were built in various sizes and orientations to correspond to the experimental factors and their settings, shown in Table 1. Minitab™ is the statistical analysis software package that was used to define the experiments and fit the response surfaces.Table 1 - Experiment Settings for each of the Build Style VariablesBuild Style Variable Layer Thickness (in.) Hatch Spacing (in.) Z-Height (in.) Part VolumeLow Value (·1) 0.004" 0.004" 0.5" 0.75 in3Middle Value (0) 0.006" 0.012" 1.0" 1.5 in3Hi2h Value (+1) 0.008" 0.020" 1.5" 2.25 in3Accuracy Prediction The factors used to fonnulate the accuracy response surface are: part height, layer thickness, and hatch spacing. Part height was chosen as a factor because the SLA is less accurate in the z-dimension than in the x and y-dimensions. Layer thickness affects build quantization errors, which detract from accuracy.

Part height was chosen as a factor because the SLA is less accurate in the z-dimension than in the x and y-dimensions. Layer thickness affects build quantization errors, which detract from accuracy. Hatch spacing has a dramatic affect on post cure shrinkage which in turn greatly affects accuracy. When the hatch spacing is less dense more resin is cured after the build process and this causes shrinkage.The unrefined response surface for accuracy is shown in Eqn. 1. The fit parameters are: S =380.0 and R-Sq =92.1 %.Accuracy =-12.6E6(H)(LT) + 11125(Z)(LT) + 48344(Z)(H) + 96.2E6(LT/ + 3470000(H/ + 577(Z/- 703617(LT) - 41629(H) - 321(Z) + 2665(1)Build Time Prediction The Build Time Estimator (BTE) program [11] was used to obtain data for the build time response surface, rather than building each parts individually. All four experimental factors from Table 1 were used.

All four experimental factors from Table 1 were used. The BTE program is usually accurate within 3 percent.The unrefined build time response equation is shown in Eqn. 2. The statistical fit parametersare: S = 48.61, R-Sq = 95.4%.Build Time =117(V)(Z) - 203(H)(Z) - 5135(H)(V) - 33813(LT)(Z) + 1125(LT)(V) -1910000(LT)(H) + 95(Zl + 42(Vl + 693238(H/ + 9470000(LTf + 33(Z) - 81(V) - 4479(H) - 70138(LT) + 301(2)where: V = part volume (in3 thickness (in).), Z = part height (in), H = hatch spacing (in), and LT = layer630The resulting equations (Eqns. 1 and 2) are used to define the build style aspiration space which determines the performance of a particular combination of build style variables.

1 and 2) are used to define the build style aspiration space which determines the performance of a particular combination of build style variables. These equations can be refined to eliminate some statistically insignificant variables, however this could reduce the accuracy of the equation with minimal benefit to computational time.Surface Finish Prediction The theoretical method for determining surface finish uses equations based on layer thickness and the surface angle. The surface angles can be determined from the STL file that describes the part. In order to predict the surface finish for a given STL file, a computer routine to estimate surface finish is used. This routine takes uses the STL flie as the input. From this information, the routine can predict the relative surface finish of several different build orientations. The equation used to predict surface finish for one facet is shown in Eqn.

From this information, the routine can predict the relative surface finish of several different build orientations. The equation used to predict surface finish for one facet is shown in Eqn. 3 and was calibrated using experimental dataPRe =(2 * cos(e) * sinee) * 937) + 3.5*e + 48(3)The predicted surface finish will be weighted by facet area, compromise DSP formulation.then normalized for use in the3.3 Compromise DSP FormulationAs mentioned in Section 2, the compromise DSP is a design decision making formulation. Values for a set of design variables are chosen to meet a set of goals as well as possible, subject to constraints. For our problem, the design variables are the build style variables, part orientation, layer thickness, and hatch spacing. Goals include build time, accuracy, and surface finish.

For our problem, the design variables are the build style variables, part orientation, layer thickness, and hatch spacing. Goals include build time, accuracy, and surface finish. The constraints for build style decision making apply to all builds; they arise due to the SLA machine limitations and manifest themselves as bounds on the design variables.For a compromise DSP, goals are formulated using Eqn. 4. In this formulation, the objective is to bring each of the goals to the minimum achievable value for that goal. That is, it is desirable to have the numbers quantifying build time, accuracy, and surface finish as close to the minimum as possible. This requires estimating the minimum (best) and the maximum (worst) possible values for each of the system goals. This normalizes the goals to ensure that no single goal dominates the solution process. The closer the goal value is to 0, the smaller the deviation variable will be.

This normalizes the goals to ensure that no single goal dominates the solution process. The closer the goal value is to 0, the smaller the deviation variable will be. For this formulation the deviation variable di' will always be zero because it is impossible to underachieve a minimum. Thus, the only deviation variable under consideration is di+.A(X)-A,min(X) + d.- -d.+ =0 Amax (X) - Ai . (X)II,mIDwhere: A(X) =The achievement function, Amin(X) =the minimum (best) possible achievement, Ai,max(X) = the maximum (worst) possible achievement. The minimum and maximum values are dependent on the part being built. The equations for A(X) were given as Eqns. 1, 2, 3 in the previous subsection. X represents the vector of build variables.(4)A deviation function measures the distance from a point in the feasible design space to the the preemptive formulation which uses goals.

X represents the vector of build variables.(4)A deviation function measures the distance from a point in the feasible design space to the the preemptive formulation which uses goals. Deviation functions can be formulated in two ways: goal priorities, and the Archimedian formulation which uses goal weights (importance values). Using the compromise DSP, the mathematical formulation of the build style decision problem for the circuit breaker handle in Section 4 is shown in Figure 1.is usually accomplished using a multiobjective optimization code such as DSIDES·[6]. However, since the build style design space is relatively small, an exhaustive search (i.e., every possible combination of build style variables) can be performed to yield a solution fairly quickly. With this method, the true minimum value of the deviation is found in about two seconds.

With this method, the true minimum value of the deviation is found in about two seconds. The ranges and number of values for each design variable are listed here along with the number of points in parentheses: Layer Thickness 0.002" 0.010" (17), Hatch Spacing 0.004" - 0.020" (17), and Part Height 0" - 10" (3). Height isSolving compromise DSP problems631dependent on three different build orientations, defined by the dimensions of a bounding box, but could be increased to many more orientations to account for support structures or other build goals not considered in this paper. Thus, the number of possible points to evaluate for any part is 17 x 17 x 3 =867.Given: STL file of circuit breaker handle. Goal importances.

Goal importances. Find: Values of system variables: Xl = layer thickness, X2 = hatch spacing, X3 = part orientationValues of the deviation variables associated with the goals: d l+ - build time, d2+ - accuracy, d3+ - surface finishSatisfy: System goals: Al (X) -116.8 - d+ = 01006.6-116.8 ~(X)-2088 d+ =0 8725-2088 ~(X) - 251.3 _ d+ =0 523.2 - 251.3 Bounds: LayerThickness 0.002" <= Xl <= 0.010" 0.004" <= X2 <= 0.020" 0" <= X3 <= 10"I23Hatch Spacing Build HeightBuild TimeAccuracySurface FinishMinimize: the Deviation Function:Z =[d3+' O.5(dt + d2+)] (Scenario 1),Z =1: Wi dt (Scenarios 2,3)Figure 1 - Compromise DSP Build Style Problem Formulation.The exhaustive search method has been embodied in a software tool called CABSS (Computer-Aided Build Style Support) and has been validated using several examples, one of which is included in the next section.4.

A TEST CASE: CIRCUIT BREAKER HANDLEThis section shows some results of applying the compromise DSP to generate build styles for different solution scenarios. The part considered is a circuit breaker handle manufactured by the Siemens corporation. This part was chosen for its geometric complexity. The breaker handle was scaled to twice its normal size so that it would lie within the bounds of the eXPerimental design space that was used to create the response surface models. When using empirically derived equations, it is important to only interpolate within the bounds of the eXPeriment space, and not extrapolate beyond these boundaries. The wireframe and STL models of the handle are shown in Figure 2. Its overall dimensions are 1.45" x 1.62" x 2.86" (X x Y x Z).4.1 Solution ScenariosThe problem of determining the build style will be solved using three different scenarios for Scenario 1 emphasizes surface finish using a prioritizing and weighting the build goals.

preemptive formulation of the deviation function. Scenario 2 places a high importance on accuracy. Scenario 3 places a medium emphasis on build time, followed by accuracy. Numerical weighting values for these scenarios are shown below. These scenarios will be used to determine how the build style variables change with respect to changing build goals.Goal Surface Finish Build Time AccuracyScenario 1 Levell, Weight 1.0 Level 2, Weight 0.5 Level 2, Weight 0.5Scenario 2 0.1 0.2 0.7Scenario 3 0.2 0.5 0.36324.2 ResultsThe STL fIle of the handle was entered into CABSS where volume, surface area, and theoretical surface fInish for three orientations were calculated. After entering the build scenarios (above), CABSS exhaustively searches the design space defmed by goal equations presented in Section 3. The results of this search are presented in Table 2.Figure 2 • Siemens circuit breaker handle.

The results of this search are presented in Table 2.Figure 2 • Siemens circuit breaker handle. Wire frame and STL renderingsTable 2 • Results from Exhaustive Search for Circuit Breaker ExampleScenario #1#2#3Build Style Variables Layer Thickness =0.0055" Hatch Spacing =0.013" Height =1.45" Layer Thickness =0.0045" Hatch Spacing =0.007" Height =1.45" Layer Thickness =0.006" Hatch Spacing =0.015" Height =1.45"Build Goals Build Time =301.28 Average Inaccuracy =2597.21 Surface Finish =251.77 Build Time =403.7 Average Inaccuracy =2239.48 Surface Finish =251.77 Build Time =271.74 Average Inaccuracy =2824.78 Surface Finish =251.77Deviation Functionlevel 1 =0.0 level 2 =0.061level 1 = 0.036level I =0.046From knowledge of SLA operation, build time decreases as layer thickness and hatch spacing increase. However, part inaccuracy increases as layer thickness and hatch spacing increase.

However, part inaccuracy increases as layer thickness and hatch spacing increase. Therefore, it is reasonable to assume· that as the importance of accuracy increases, layer thickness Increasing the importance of build time should have the and hatch spacing should decrease. opposite effect on these values. As can be. seen from Table 2, layer thickness and hatch spacing In Scenario 2, when attain their largest values in Scenario 3, when build time is most important. accuracy is most important, layer thickness and hatch sp~ingtake on their smallest values.In order to validate the results obtained; the handle should be built with the ''best'' build style variable values. At present, only the build time goal has been tested. The BTE program provided the baseline build time measurements. Comparison of BTE and CABSS build times shows an error of less than 2 percent for Scenarios 1. and 3 and an error of 5 percent for Scenario 2.6335.

Comparison of BTE and CABSS build times shows an error of less than 2 percent for Scenarios 1. and 3 and an error of 5 percent for Scenario 2.6335. CONCLUSIONSSeveral conclusions can be drawn from this research. The main conclusion is that it is indeed possible to provide decision support for stereolithography build styles using the compromise nsp. This is done by formulating and solving a nsp with a computer tool. Several specific conclusions are used to support this main conclusion.The selected build style variables have a significant impact on SLA build performance: build orientation (part height and surface angles), cross sectional layer thickness, and laser hatch density. This research verified the expected behavior of these variables.Build performance can be predicted by constructing response surface models of the build goals as functions of the relevant build style variables and parameters.

These response surface models apply to parts of a wide range of geometric complexities and sizes. This conclusion is based on observing the test case results and comparing them with expected SLA build behaviors. However, the response surface models should only be applied to parts whose parameters fit within the experimental design space on which the response surface is based.Even though this method for decision support is viable, more work needs to be done to improve the results. This includes creating more accurate response surface models that encompass a wider range of part parameters (volume arid height). ACKNOWLEDGEMENTSWe gratefully acknowledge the support from ARPA TRP grant #2187 and from the RPM!member companies.

ACKNOWLEDGEMENTSWe gratefully acknowledge the support from ARPA TRP grant #2187 and from the RPM!member companies. REFERENCES1Frank, D and Fadel, G, "Expert System Based Selection of the Preferred Direction of Build for Rapid Prototyping Processes," Clemson University report, 1993.2 Wodziak, J R, Fadel, G M, Kirschman, C, "A Genetic Algorithm for Optimizing Multiple Part Placement to3Reduce Build Time," Clemson Intelligent Design Systems Laboratory, Clemson University, 1994. Sabourin, E, Houser, S A, BS(jhn, J H, "Adaptive Slicing using Stepwise Uniform Refinement," Rapid Prototyping Journal, 2(4):20-26, 1996.4 Hur, J and Lee, K,for Automatic Support Structure Generation in Layered Manufacturing," Proceedings ASME Computers in Engineering Conference, paper #96-DETC/CIE-1324, 1996.

5 McClurkin, J E and Rosen, D W, "A Build Style Selection Method for Stereolithography," Proceedings 1997"Efficient AlgorithmsNASUG Conference, Orlando, February 16-20, 1997.6 Mistree, F, Hughes, 0 F, Bras, B A, "The Compromise Decision Support Problem and Adaptive Linear Pro gramming Algorithm," Structural Optimization: Status and Promise, AIAA, Washington, DC, pp. 247-289, 1993.7 Mistree, F, Hughes, 0 F, Phuoc, H B, "An Optimization Method for the Design of Large, Highly Constrained8Complex Systems," Engineering Optimization, 5(3):141-144, 1981. Tata, K and Flynn, D, "Quantification of Down Facing Z Error and Associated Problems," Proceedings 1996 NASUG Conference, San Diego, March 11-13, 1996.9 Maestro 1.8 Users Manual, 3D Systems Inc., Valencia, CA, 1996.

ild Time Estimations for Large Scale ModellingG. Mensing and 1. Gibson University of Hong KongAbstract Achieving speedy results in model making is very much desired if not a necessity in ahnost any manufacturing industry. There is no doubt that rapid prototyping contributes to this process. It is generally considered that when compared to conventional machining techniques like nlilling, the current rapid prototyping systems appear to be much faster. This is certainly true for complex, slnall objects. I-Iowever, this is not alwaysa,pplicable to simple, large and bulky parts. There are a number of projects and systems concentrating on the fabrication of large models. Work is being carried out at the University ofHong Kong, using milling. along with slicing technology. This.report compares some of the rapid prototyping systems witl1milling.

Work is being carried out at the University ofHong Kong, using milling. along with slicing technology. This.report compares some of the rapid prototyping systems witl1milling. Milling is an established technology and recent developments in materials and nlachines used in Inilling nlake it a good alternative to rapid prototyping when itcomes to largesyale nl0delling.Introduction Industry has often expressed the need for prototypes larger than current rapid prototyping technology is capable of. A number of approaches [1] [21have been developed to address this. The system at the University of Hong Kongis similar to those developeclatFord Research [3]. This report follows as a result of the proposed system for large scale rapid prototyping [4]. Milling, combined withlayer based technology results in fasterprototypingfor large, bulky objects. Milling is also more appropriate for tooling applications because of the variety of Inaterials that can .be used.

Milling is also more appropriate for tooling applications because of the variety of Inaterials that can .be used. Some simple objects of variable sizes are discussed with the view to understanding how rapid prototyping systems would deal with making models thatare larger than they are currently clapabIe of.. Extrapolation is used to calculate the times needed for the largest models since these.are too big for the current systems. The systems are compared using build time. Conlpanies are willing to pay more in the short term as long as there is a considerable saving in time. This makes the conlpany nlore conlpetitiveand .can save money in the long run. However rapid prototyping Inay not save time when building large models compared with conventional technology.1. The rapid prototyping systems First SOine rapid prototyping systems are discussed. These are the systems that wereusedinthe test.

The rapid prototyping systems First SOine rapid prototyping systems are discussed. These are the systems that wereusedinthe test. Each system displays characteristics that may make it difficult to construct thenl so they are capable of making larger models.1.1.1. StereoLithography - 3D-Systems StereoLithography [5] [6] uses photo-reactive polymers,which react to this,the system utilises a precisely directed laser beam.• When resin is struck by oJ"Uc.l.~""'~'""~~"it solidifies (polymerises). By scanning UV light onto the surface of the polymer according to the design, a layer (slice) of a model. is created. To make the next and subsequent layers, the object is lowered into the vat filled with resin. This way new resin flows Qver the. object. Next, this new layer of resin is solidified. The process continues until the object is finished. This process is accurate, but.very limited in the variety of materials.

object. Next, this new layer of resin is solidified. The process continues until the object is finished. This process is accurate, but.very limited in the variety of materials. It is possible though to use the models to manufacture moulds for casting.1.1.2.iModelMaker __ Sanders Prototype Sanders1'7] developed the Model Maker, aplotting system where a liquid-to-solid inkjet plotter with a §eparate Z-axis inputis used. An inkjet subsystem rides on a precision X/Y drive, carriage and deposits both thermoplastic and wax materials on the build substrate. The X/Y drive carriage also energises a flatbed milling subsystem for maintaining precise Z-axis dimensioning of the model by milling off the excess vertical height of the current build layer. A support material is used to support..overhangs andcavitiesinthemodel during the build. Droplets are placed upon the build substrate to within 0.00025 inches (0.007 rom) in the X and Y directions.

A support material is used to support..overhangs andcavitiesinthemodel during the build. Droplets are placed upon the build substrate to within 0.00025 inches (0.007 rom) in the X and Y directions. The droplets stick to each other during the liquid-to-solid phase tran~ition to form a uniform mass. After solidification the millingof the layers immediately follows the deposition cycle. It is a slow but accurate process and only thermoplastic materials can be used.1.1.3.•. Fused Deposition Modelling - Stratasys With FusedPeposition Modelling.or FDM[8], a temperature-controlled head extrudes thermoplastic material layer by layer and thus creating a model. The CAD-model is sliced and, if necessary, supports are created similar to the StereoLithography process. Path data is then downloaded to the FDM system. The system operates in the X, Y and Z-axes. In effect, it draws the model one layer at a time.

Path data is then downloaded to the FDM system. The system operates in the X, Y and Z-axes. In effect, it draws the model one layer at a time. Thermoplastic modelling material, a wire of 0.070 inches (.178cm) in diatTIeter feeds into the temperature-controlled FDM extrusion head, where it is heated to a semi-liquid state. The head extrudes and deposits the material in ultra-thin layers (0.05 to 0.762 mm) onto a fixtureless base. The head directs the material into place with precision. The material solidifies, laminating to the preceding layer.1.1.4. Selective Laser Sintering - DTM The Sinterstation, from DTM [9] works with powder, which is selectively sintered by means of a CO2 laser. A roller spreads powder on a bed. A laser selectively sinters this layer. The lTIodel is indexed down. This process continues until the model is completed. An advantage of this process is that it does not need supports.

A laser selectively sinters this layer. The lTIodel is indexed down. This process continues until the model is completed. An advantage of this process is that it does not need supports. DTM is in the process of developing a wider variety of materials, and although metal powder can be used it is not quite good enough for long'term tooling.Laminated Object Manufacturing -KiraThe machine from Kira [10], Japan, is a Rapid Prototyping Machine suitable for installation in an office environment. The machine uses paper (wood) as material and the cutting process is not carried out with a laser as used by Helisys [11], but is realised mechanically by means of a cutting plotter. A laser printer is used to print the outline of the model onto a sheet of paper. This paper is344then bonded to any previous layers. The Kira then cuts the outline ofthe shape of this layer. Next a new sheet of paper is added on t()p.ofthe.former, bonded and thencut.

This paper is344then bonded to any previous layers. The Kira then cuts the outline ofthe shape of this layer. Next a new sheet of paper is added on t()p.ofthe.former, bonded and thencut. This continues untilthe model is finished.1.1.6. DeskProto [milling] - Delft Spline Systems DeskProto [12], a software package, reads the STL-file and displays its contents. Next some milling parameters such as the type of cutting tool, required accuracy, etc. are entered. DeskProto then automatically calculates the milling paths. These milling paths. are sent to the desktop NC milling machine, which produces the object. This process can be done with a wide variety of materials.1.2. Current Rapid Prototyping Systems on a large scaleAn obvious technique for producing large scale prototypes would be to increase the size of the current systems. This .might not be a trivial process.

This .might not be a trivial process. Increasing the size of all the systems .would be costly, requiring much more attention to system stability and structure. With StereoLithography, for instance, system designers would have to consider the problem of maintaining the stability of the resin over a large surface.· area. Laser based systems would require much more powerful lasers and larger surface areas would also result in significant problems with the optics. A number ofthe systems use plotting mechanisms (e.g. Sanders, FDM). Large prototypes would require a significant amount of travel for these mechanisms which would result in long build times. Kira, Helisys and milling rely more on the surface area than the volume of a part. This implies that fabrication of large bulky objects using these techniques would require less energy than purely additive processes.CubeCylinderHollow Cube2.

This implies that fabrication of large bulky objects using these techniques would require less energy than purely additive processes.CubeCylinderHollow Cube2. The test objectsHol19wSphere [Cut-outfOt.·reference]For this experiment only simple objects are chosen. With basic forms, it is possible to get an idea on how much time a more complex. form would need, since most shapes are combinations of the basic forms. The data obtained was primarily from build time estimators, simple geometry therefore makes it easier to produce an estimation without significant amounts of processor, user or actual build time. Users of the different technologies were. asked to provideestima.tionsJorbuilding the defined geometries, figure 1.

Users of the different technologies were. asked to provideestima.tionsJorbuilding the defined geometries, figure 1. Every objectis available in four different sizes, this is to find the relationship between size and speed for the current layer based systems and milling and also to provide a means for extrapolation.Figure 1345The objects were 1uade in a vertical plane. This set-up is not necessarily the best for optimising the rapid prototyping system parameters. Sonle rapid prototyping systems have the option to change solid structures into hollow or honeycomb structures. This has not been done with the nl0dels for this report. There is a solid and a hollow version of each model in this report, which may indicate the potential tinlesaving when using hollow build styles.The experimental build data3.1.1. 3D-Systems - SLA-500, SLA-250 The build times were calculated with estimation software.

3D-Systems - SLA-500, SLA-250 The build times were calculated with estimation software. While laser scan time depends mostly on the laser power, recoating time goes down significantly with thicker slices and slualler parts. The layer thickness that was used is 0.25 mm and STAR-Weave was used.Sanders -Model MakerIn the Model Maker II a 0.052 lUlU slice thickness was used for all but the largest nl0dels used in the experiment, where a 0.13 nlm slice thickness was used. The build parameters are set by using what is called a 'Configuration'. This includes not only the slice thickness, but also the thickness of the line, velocities, line spacing, etc. 'Configuration' #470 ("Fast Regular") and #395 ("Concept Model") were used for all the builds. For the cubes and cylinders, an option to generate no support stl'UctltreS was also used, since none is needed. For the other parmueters the system defaults were used.3.1.3.

For the cubes and cylinders, an option to generate no support stl'UctltreS was also used, since none is needed. For the other parmueters the system defaults were used.3.1.3. Stratasys - FDM-2000 The FDM-2000 worked with a layer thickness 01'0.25 1um. The object was not actually built. The build times were calculated with the build time estimator software that comes with the system, . using the default settings.3.1.4 DTM - Sinterstation 2000 The layer thickness used for the sinterstation was 0.25 lum. For the cubes and the cylinders only a 1 mm layer was made in Trueforrn and from this the total build tilne was calculated. Only half of the spheres were built because of sYlumetry. From this the build tinles were calculated.3.1.5. Kira - Laminated Object Manufacturing The objects were only partly built, for the cylinder and cube only Inlm layers were fabricated. The paper is 0.085 mm. From these times the total build times were calculated.

The paper is 0.085 mm. From these times the total build times were calculated. The build tilues for the hollow objects is on the low side since minimal cross-hatching was used for the calculation. This may result in difficulties during the decubing process.3.1.6. Delft Spline Systems - DeskProto The milling estimates are very rough estimates, as the actual build times are very much dependant on the capabilities of the CNC milling machine and material used. For the larger models a lot of practical problems can be expected. For instance the steep vertical walls will need a very long cutter or a five-axis machine. These problems are not taken into account: assuming them to be solved.346The time Inentioned in table 1 is for an accurate n10del in tooling board (a wood-like 111aterial). The build tin1es for a cube (solid and hollow) and a cylinder (solid and hollow) are considered the smne.

The build tin1es for a cube (solid and hollow) and a cylinder (solid and hollow) are considered the smne. This is done because the travelling tin1e of the n1ill head is about the smne for soft n1aterials. Build tilnes are also very dependent on the level of details: The more con1plex the 1110deL the slnaller the .tool required, therefore the longer the build tilne. For the basic n10dels in this test a large tool is considered: the larger the InodeL the larger the tool. The hollow sphere is Inade in two parts and later joined together.The ResultsIT:1. r..

With the objects being basic shapes, the data for the cube, cylinder and the sphere are relatively the same, which therefore give comparable graphs. That is why only the graphs for the cube and the hollow cube are shown. Graph 3 shows a comparison between outline scanning (milling, Kira, etc.) and volume scanning (StereoLithography, FDM, SLS. etc.). Volume scanning systems represented in the graph employ a scanning width of 0.3 mm. This figure is quite high for StereoLithography and selective laser sintering. These systems normally use a scanning width of about 0.2 mm. The FDM-2000 uses a scanning width of about 0.3 mm. but can use wider scanning widths. These results are likely to be prone to error but even considering an error of 20% it is still clear that outline scanning systems are more competitive. A 20% error is acceptable based on the generally considered performance of the build time estimators and the true build time of the Sinterstation and the Kira.

A 20% error is acceptable based on the generally considered performance of the build time estimators and the true build time of the Sinterstation and the Kira. For the smaller objects the time difference between volume scanning systems and outline scanning systems is not that different, however for the larger parts the difference is significant. Rapid prototyping systems like ZCorp and Stratoconception are not mentioned. This is because data was not available. It is understood that the ZCorpmachine in particular would provide a significant reduction in build time and should be the subject of further studies.Graph 3. Outline Scanning versus Volume Seaming10000000'i/Outline Scanning/', /E 1000000 .§. CI> (.) c .t9 tI> C 0) c 'c c co (.)en100000.,j,"">->.,.---.-)¥;-----------....

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperBUILDING ZR-BASED METALLIC GLASS PART ON Ti-6AL-4V SUBSTRATE BY LASER-FOIL-PRINTING ADDITIVE MANUFACTURINGYingqi Li, Yiyu Shen, Ming C. Leu, and Hai-Lung TsaiDepartment of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractThrough using Zr intermediate layers, Zr52.5Ti5Al10Ni14.6Cu17.9 metallic glass (MG) parts are successfully built on Ti-6Al-4V substrate by laser-foil-printing (LFP) additive manufacturing technology in which MG foils are laser welded layer-by-layer onto the substrate. The printed MG part is free of porosity, cracking and crystallization, and its properties are very similar to the original MG material.

The printed MG part is free of porosity, cracking and crystallization, and its properties are very similar to the original MG material. The Zr intermediate layers are aimed at preventing direct interaction between the first layer of MG foil and the Ti substrate; otherwise, the welded MG foils would peel off from the substrate due to the formation of brittle intermetallic compounds. With the use of Zr intermediate layers, the bonding strength of the printed MG part and the Ti substrate can reach 758 MPa owing to the formation of α-Zr phase.Keywords: Metallic glass; Titanium; Laser welding; Embrittlement; Additive manufacturing.IntroductionThe amorphous atomic structure and concomitant lack of grain boundaries or slip planes endow metallic glasses (MGs), i.e., amorphous alloys, with many superior properties, such as high strength, corrosion resistance, wear resistance, etc., compared with conventional crystalline metals [1-3].

However, current applications of MGs are limited by small dimensions and simple geometries (rod, foil, thin plate, etc.) because fast cooling is necessary to produce fully amorphous structures [2].Laser-based additive manufacturing (AM) technology is a promising method to fabricate large and complex MG parts owing to the very high heating and cooling rates of laser processing. Recently, a variety of three-dimensional MG parts were successfully built without crystallization using a newly developed laser-foil-printing (LFP) AM technology [4]. In this technology, MG foils are laser welded layer-by-layer upon the substrate to produce MG parts. Since metal foils, instead of powders, are used as feedstock, the LFP technology shows its unique advantages over powder-based AM methods on the fabrication of MG structures, such as easy to fabricate full- density part, higher cooling rate at the same laser processing condition, etc.

The LFP technology can expand MG products to 3D freeform geometries with large dimensions.Normally, a large, thick metal plate is used as the foundation substrate in LFP technology upon which additional materials are gradually added layer-by-layer. Since it is difficult or even impossible to obtain large, thick as-cast MG plates, conventional crystalline metals have to be used771as substrates. Thus, a good bonding between the MG foil and the crystalline substrate is crucial to the success of MG structures manufactured by LFP; otherwise, the welded MG foils would peel off from the substrate before the desired structure has been built because of the accumulation of residual stresses.Many engineering applications require structures or parts with different properties in different locations. For example, a structure may need corrosion or wear resistance in one area while needs high toughness in another area.

For example, a structure may need corrosion or wear resistance in one area while needs high toughness in another area. Besides properties, cost of the structure may also need to be taken into consideration. But rarely a single material can meet all these requirements. The LFP technology provides unique opportunity for printing a structure consisting of multiple materials. Theoretically, any metal which can be laser welded is a candidate material for LFP. Resultant property changes may be obtained through welding various metal foils at different desired layers or locations during LFP to meet various application needs. MG has superior corrosion resistance and wear resistance, but its intrinsic fracture toughness is usually insufficient to allow structures entirely made of MGs to obtain optimum performance [5]. In addition, to date many commercially available MGs are made from expensive metals with high-purity, e.g., high-purity Zr, resulting in very costly MGs.

In addition, to date many commercially available MGs are made from expensive metals with high-purity, e.g., high-purity Zr, resulting in very costly MGs. These factors also limit MGs’ application. In many applications, there is no need to fabricate an entire MG structure; instead, most portions of the structure may be printed using cost-competitive conventional metals with favorable toughness, and only components that need to serve in extreme environment (e.g., corrosive, wear) are printed with MG. The ability to print MG parts on conventional crystalline metals, e.g. titanium alloys, stainless steels, etc., would extend the engineering applications of MG, and offer more flexibility in structure design and performance.Laser welding of MG to crystalline metals is therefore required, but is very challenging due to the large differences in physical and chemical properties which may be present between them.

Laser welding of MG to the same MG has been studied in butt-joint or bead-on-plate form [6-9]. However, there is little or no publication on laser welding of MG to crystalline metals. In this study, Zr52.5Ti5Al10Ni14.6Cu17.9 MG parts were built on Ti-6Al-4V (Ti64) substrates. Our previous results show that the welded Zr-based MG foils would peel off from the Ti64 substrate when directly building an MG part on a Ti64 substrate due to the formation of hard and brittle intermetallics [10]. A feasible way to reduce the brittleness of the dissimilar bonding is to use an intermediate layer to separate the two dissimilar materials. The intermediate layer should be highly soluble with both materials. Zr has similar chemical properties to Ti, and Ti-Zr system can form hcp solid solution over the complete composition range [11].

The intermediate layer should be highly soluble with both materials. Zr has similar chemical properties to Ti, and Ti-Zr system can form hcp solid solution over the complete composition range [11]. On the other hand, Zr is the main element of the selected MG, and Zr-based MG parts can be built on Zr substrates by LFP without peeling off due to the formation of α-Zr phase in the MG-Zr interface [10]. Thus, pure Zr foil was chosen as the intermediate layer to improve the bond between Zr-based MG part and Ti64 substrate in this study. The properties of the printed MG part will be discussed first, and then this paper will focus on the study of the dissimilar bonding between the MG part and the Ti64 substrate.ExperimentalSince this paper focuses on the bonding between Ti64 alloy and MG, 12-mm-thick Ti64 plates, instead of 3D printed Ti64 parts, were used as the substrates in this experiment to shorten the fabrication time.

The thickness of a layer of MG foil with nominal chemical composition of772Zr52.5Ti5Al10Ni14.6Cu17.9 (at. %) (Liquidmetal®, LM105) was approximately 200 μm. 200-μm- thick pure Zr foil (Zr 702) was used as the intermediate layer. Before laser processing, the substrates and foils were slightly grinded using abrasive papers to remove surface oxides, and then cleaned using ethanol.The LFP system developed at the Laser-Based Manufacturing Lab, Missouri S&T, schematically shown in Fig. 1, was used in this experiment. This system mainly consists of two lasers: a continuous wave fiber laser (IPG YLP-1000) with 1070 nm wavelength for foil-welding and a pulsed UV laser (Coherent AVIA-355X) with 355 nm wavelength and 30 ns pulse duration for foil-cutting. In this study, the welding laser beam entered a scanner (SCANLAB) and then was focused on the foil surface, and the beam diameter (D) at the focus was about 170 µm.

In this study, the welding laser beam entered a scanner (SCANLAB) and then was focused on the foil surface, and the beam diameter (D) at the focus was about 170 µm. The substrate was mounted on a computer controlled three-axis motion stage (Aerotech). A layer of foil was placed and fixed on the substrate via laser spot welding using a mold-plate with a predetermined hole-pattern (see Fig. 2(a)). Then, the scanner scanned multi-passes to weld the foil onto the substrate, as shown in Fig. 2(b) and (d). After the welding process was done, the motion stage shifted the workpiece to the location under the UV laser and the foil was cut by the UV laser along the contour of the welded region to remove the redundant foil (see Fig. 2(c)). Another layer of fresh foil was placed on the top of the workpiece for the process of next layer. The above processes were repeated until the desired part was built.

2(c)). Another layer of fresh foil was placed on the top of the workpiece for the process of next layer. The above processes were repeated until the desired part was built. All samples were processed inside a chamber with Ar shielding gas flow.Figure 1 The LFP system developed at the Laser-Based Manufacturing Lab, Missouri S&T. Reprinted from Ref. [12].773Figure 2 (a-c) Schematic of the LFP process; (d) schematic view of the transverse cross-section after the first layer of foil is welded onto the substrate. Δd is the hatch space, and δ indicates the penetration. Reprinted from Ref. [10].Two cases were compared in this paper, i.e., without any intermediate layer (Case 1) and with Zr intermediate layers (Case 2). For Case 1, MG foils were directly welded onto the Ti64 substrate, and only ten layers of MG foil were welded since the welded foils peeled off from the substrate before the desired part was built. For Case 2, the first and second layers were both Zr foils.

For Case 2, the first and second layers were both Zr foils. After that, 25 layers of MG foil were further welded onto the workpiece to construct a 5-mm-thick MG part with a base area of 15 mm× 20 mm (see Fig. 3). To reduce heat accumulation, the time lapse between two adjacent scan passes was 10 s. The used laser power (P), laser scanning speed (v) and hatch space (Δd) for each layer of welding are listed in Table 1.

The used P and v values were selected after parametric study to ensure continuous and stable welding passes and small penetration depth at the same time.Table 1 Parameters for each layer of weldingLayer1st layer of Zr foil (if any) 2nd layer of Zr foil (if any） 1st layer of MG foil 2nd-25th layer of MG foilLaser power (P) (W) 500 500 500 500Laser scanning speed (v) (mm/s) 500 450 570 650Hatch space (Δd) (mm) 0.12 0.12 0.12 0.14774Figure 3 (a) Schematic of the printed MG part built on Ti64 substrate with intermediate layers; (b) schematic of the bonding between MG and Ti64 when using two layers of Zr foil as intermediate layers; (c) dimensions of the miniature tensile test specimen.After the MG part was built, miniature tensile test specimens were cut from the substrate and the printed MG part by wire electric discharge machining (EDM), to test the bonding strength between the Ti64 substrate and the MG part, as schematically shown in Fig. 3(a).

3(a). The dimensions of the miniature tensile test specimens are shown in Fig. 3(c), and the thickness was 1 mm. The tensile test was conducted on an INSTRON-5969 testing machine. Five specimens were tested, and the mean value was reported. The tensile strength of the printed MG part was also tested using the same method. The microstructure and chemical composition were analyzed using X-ray diffraction (XRD), optical microscope (OM), and scanning electron microscope (SEM) equipped with energy dispersive spectroscopy (EDS). The glass transition temperature and crystallization temperature of the MG part were measured by differential scanning calorimeter (DSC). The metallographic specimens for OM and SEM analysis were prepared following the general procedure of metallurgical sample preparation.Results and DiscussionFigure 4(a) shows the OM image of the transverse (perpendicular to the laser scanning direction) cross-section of the printed MG part.

This cross-section was etched using a solution of 100 ml H2O, 2 ml H2O2 and 0.1 ml HF which can reveal the crystallization of LM105 MG occurred during laser scanning [4]. It can be seen that the cross-section is featureless; porosity, cracking and crystallization were not observed. The XRD result of the printed MG part, which was conducted on the longitudinal (parallel with the laser scanning direction) cross-section of the middle plane, also confirms that the printed MG part is fully amorphous within the detection limit of XRD (~ 7% [13]) (see Fig. 4(b)). DSC was used to compare the glass transition temperature (Tg) and crystallization temperature (Tx) of the printed MG part and the original MG at a heating rate of 20 °C/min, and the results are shown in Fig. 4(c). For the printed MG part, Tg ≈ 397 °C, Tx1 ≈ 475 °C and Tx2 ≈ 482 °C. These values are very close to those of the original MG, in which Tg ≈394 °C, Tx1 ≈ 470 °C and Tx2 ≈ 478 °C.

4(c). For the printed MG part, Tg ≈ 397 °C, Tx1 ≈ 475 °C and Tx2 ≈ 482 °C. These values are very close to those of the original MG, in which Tg ≈394 °C, Tx1 ≈ 470 °C and Tx2 ≈ 478 °C. Tensile tests were conducted to investigate the mechanical properties of the printed MG part. For comparison, the original LM105 MG with a thickness of 750 μm was also tested. The tensile test was conducted along the laser scanning direction, and four samples were tested and the average value was reported. The original MG has a yield strength of ~ 862 MPa and an ultimate strength of ~ 1533 MPa with an elongation-to-failure of ~ 5.8%. The printed MG part exhibits higher yield strength (~ 1138 MPa) and ultimate strength775(~ 1758 MPa), but slightly lower elongation-to-failure (~ 4.9%).

The printed MG part exhibits higher yield strength (~ 1138 MPa) and ultimate strength775(~ 1758 MPa), but slightly lower elongation-to-failure (~ 4.9%). The above observations show that the printed MG part is extremely similar to the original MG.Figure 4 The printed MG part: (a) OM image of the transvers cross-section; (b) XRD pattern conducted on the longitudinal cross-section of the middle plane; (c) DSC profile; (d) tensile stress-strain curve. For comparison, the data for the original MG were also included.Figure 5 shows the bonding area between the MG part and the Ti64 substrate without intermediate layer (subfigure (a)), or with Zr intermediate layers (subfigure (b)). When an MG part was directly built on the Ti64 substrate, the welded MG foils completely peeled off from the substrate after 10 layers of MG foil were welded, indicating a weak bonding between the MG and the Ti64, as shown in Fig. 5(a).

5(a). On the contrary, when Zr intermediate layers were used, the sample did not fracture during the part fabrication (see Fig. 5(b)). Tensile test was then performed on the joints with intermediate layers to measure the bonding strength between the MG part and the Ti64 substrate. The stress-strain curve is shown in Fig. 6(a). The average bonding strength between the MG part and the Ti64 substrate is about 758.5 MPa. The tensile test specimens fractured at the interface between the first layer of Zr foil and the Ti64 substrate, as shown in Fig. 6(b). The tested bonding strength is 20% less than the ultimate tensile strength (UTS) of Ti64 alloy (~950 MPa), but is 62% higher than the UTS of Zr 702 (~468 MPa).

6(b). The tested bonding strength is 20% less than the ultimate tensile strength (UTS) of Ti64 alloy (~950 MPa), but is 62% higher than the UTS of Zr 702 (~468 MPa). However, the specimen shows nearly no plastic deformation prior to failure (with an elongation to fracture of only ~ 0.5%), indicating a brittle fracture.776Figure 5 The bonding between the MG part and the Ti64 substrate: (a) without intermediate layer; (b) with Zr intermediate layers.Figure 6 (a) Tensile strain-stress curve of the bonding between the MG part and the Ti64 substrate with Zr intermediate layers; (b) OM image showing the fracture position of the tensile test specimen.The chemical composition and microstructure were then analyzed and compared on the bonding area between the first layer of MG foil and the Ti64 substrate for these two cases. A layer of MG foil was first directly welded onto the Ti64 substrate without any intermediate layer at P = 500 W, v = 580 mm/s, Δd = 0.25 mm.

A layer of MG foil was first directly welded onto the Ti64 substrate without any intermediate layer at P = 500 W, v = 580 mm/s, Δd = 0.25 mm. Figure 7(a) is the backscattered electron image of the transverse cross-section of the bonding area. It can be seen that the welding is a keyhole-mode, instead of conduction-mode, which is consistent with the calculated laser intensity I = P/(πD2/4) = 2.2×106 W/cm2. The depth of penetration (δ) is about 95 µm. In addition, the transverse cross- section shows a swirling structure, especially at the bottom of the welding metal, i.e., the penetration area or the MG-Ti interface. Note the interface refers to the penetration area in this paper. This swirling structure indicates the incomplete mixing between the melted Ti64 substrate and the MG foil due to rapid solidification in laser welding.

This swirling structure indicates the incomplete mixing between the melted Ti64 substrate and the MG foil due to rapid solidification in laser welding. Figure 7(b) shows the longitudinal cross-section of the bonding area, from which cracking can be easily observed at the MG-Ti interface.777Figure 7 The bonding between the first layer of MG foil and the Ti64 substrate without intermediate layer: (a) backscattered electron image of the transverse cross-section; (b) OM image of the longitudinal cross-section.Figure 8 shows the distribution of the major elements (Zr and Ti) in the MG-Ti bonding corresponding to the rectanglular area in Fig. 7(a). Table 2 summarizes the chemical composition of the tested points in Fig. 8. It can be seen that the MG-Ti interface, which has highest cracking susceptibiliy, is a Ti-rich area that has higher Ti content and less Zr than the MG foil (see Fig. 8 and Table 2).

8. It can be seen that the MG-Ti interface, which has highest cracking susceptibiliy, is a Ti-rich area that has higher Ti content and less Zr than the MG foil (see Fig. 8 and Table 2). In addition, the distribution of Ti is not uniform in the penentration area (see Fig. 8(c) and (d)), which is consistent with the swirling structure shown in Fig. 7. An addition of 2-5 at. % of Ti can improve the glass forming ability of Zr-Cu-Ni-Al amorphous alloys; however, for higher Ti content (≥ 10 at. %) and lower contents of Cu, Ni and Al (Cu + Ni + Al ≤ 35 at. %), the glass forming ability decreases and, in turn, the formation of crystalline is promoted [14]. In most of the penetration area, the Ti content is larger than 50 at. %, and Cu + Ni + Al ≤ 35 at. %, e.g., points 1-4 (see Table 2). Thus, the formation of crystalline phases is promoted.

In most of the penetration area, the Ti content is larger than 50 at. %, and Cu + Ni + Al ≤ 35 at. %, e.g., points 1-4 (see Table 2). Thus, the formation of crystalline phases is promoted. There are at least six elements (Zr, Ti, Al, Ni, Cu, and V) existing in the penetration area, and these elements are not all mutually soluble [15], which would increase the possibility of forming a large amount of intermetallic compounds. The XRD pattern, tested at the MG-Ti interface after removing the MG by mechanical grinding, shows that the microstructure is a blend of amorphous phase and intermetallics (could be ZrCu2, Ni2Ti, Ni2Zr3, Al2Zr), as shown in Fig. 9. Intermetallics are often hard and brittle, their formation may be responsible for the high cracking susceptibility in the penetration area [16].

9. Intermetallics are often hard and brittle, their formation may be responsible for the high cracking susceptibility in the penetration area [16]. On the other hand, the large difference in coefficient of thermal expansion (CTE) between LM105 MG (12.0 µm/m-°C) and Ti64 (8.6 µm/m-°C) would result in large residual stresses, which might promote the formation of cracking in the area with high cracking susceptibility.778Figure 8 (a) and (b) Element maps of Zr and Ti corresponding to the rectangular area in Fig. 7(a); (c) element distribution along the white arrow in (a); (d) magnified element map of Ti. Only the major elements are shown here.Table 2 EDS point analysis for Fig. 8 (at.

The penetration depths δ1 for welding the first layer of Zr, δ2 for welding the second layer of Zr, and δ3 for welding the first layer of MG are around 67 µm, 115 µm, and 84 µm, respectively, as shown in Fig. 10(a). This bonding consists of three interfaces, i.e., Zr-Ti interface, Zr-Zr interface and MG-Zr interface (see Fig. 10(b)). Unlike the case without intermediate layer, cracking was not observed in these three interfaces from the cross-sections.Figure 10 The bonding between the first layer of MG foil and the Ti64 substrate with Zr intermediate layers: (a) OM image of the transverse cross-section; (b) OM image of the longitudinal cross-section.Figure 11 shows the element distribution at the bottom of the first layer of Zr foil corresponding to the area “A” in Fig. 10(a). The chemical composition of the tested points on Fig. 11 is listed in Table 3.

10(a). The chemical composition of the tested points on Fig. 11 is listed in Table 3. It can be seen that the chemical composition distribution of the intermediate layer is not very uniform, caused by incomplete mixing between the melted Ti64 substrate and Zr foil, as shown in Fig. 11(a) and (b). Generally, there is ~30 at. % of Ti mixed with Zr in the intermediate layer (Fig. 11(c)), e.g., points 2-4. Figure 12 shows the chemical composition distribution at the MG-Zr interface, corresponding to the area B in Fig. 10(a). The chemical composition of the tested points in Fig. 12 is also shown in Table 3. For the second layer of Zr, the distribution of chemical composition fluctuates in the area close to the first layer of Zr, and becomes relatively uniform in the area close to the MG, as shown in Fig. 12(b) and (d). Additionally, there are ~10 at. % of Ti780and ~ 90 at. % of Zr in the area with uniform chemistry distribution, e.g., points 6-8 (see Table 3).

For comparison, the XRD pattern of the original Zr foil is also included. We can see that the microstructures at the Zr-Ti interface and at the Zr-Ti interface both are α-Zr solid solutions. However, because the solute elements and their percentages that are incorporated in the hcp crystalline lattice of α-Zr are different for the Zr-Ti and Zr-Zr interfaces, their XRD patterns are not completely the same. For the Zr-Ti interface, the solute elements are Ti, Al, and V, but the solute is mainly Ti for the Zr-Zr interface (Table 3). In addition, the solute content is higher at the Zr-Ti interface than that at the Zr-Zr interface. Thus, the peak shift and broadening of the XRD pattern is more evident for the Zr-Ti interface (see Fig. 13). For the MG-Zr interface, because of the very high content of Zr (see Fig. 12 and Table 3), α-Zr phase formed. In addition, amorphous phase and ZrCu intermetallic are also detected.

13). For the MG-Zr interface, because of the very high content of Zr (see Fig. 12 and Table 3), α-Zr phase formed. In addition, amorphous phase and ZrCu intermetallic are also detected. Because of the very low content of Cu at the MG-Zr interface (Table 3), the formation of ZrCu would not considerably contribute to the brittleness of the weldment. Compared with the intermetallics formed at the bonding between the MG and Ti64 without intermediate layer, α-Zr phase is more soft and ductile. Therefore, the improved bonding strength between the MG and Ti64 with Zr intermediate layers seems to be contributed by the very high Zr content and the resultant formation of α-Zr phase.782Figure 13 XRD patterns obtained at the MG-Zr interface, Zr-Zr interface and Zr-Ti interface.ConclusionsZr52.5Ti5Al10Ni14.6Cu17.9 metallic glass parts were successfully built upon Ti64 substrates by laser-foil-printing (LFP) additive manufacturing in this paper.

Pores, cracking and crystallization were not observed in the printed MG part, and its glass transition temperature, crystallization temperature, and tensile strength are nearly the same as the original MG material. When directly building a MG part on the Ti64 substrate, the welded MG foil peeled off from the substrate due to the formation of brittle intermetallic compounds at the MG-Ti interface caused by excessive mixing of the melted Ti substrate with the MG. When Zr intermediate layers were used, the Ti64 and the MG part had a bonding strength of ~758 MPa, and the tensile test specimen fractured at the Zr-Ti interface. This improved bonding is contributed by the very high Zr content and the resultant formation of α-Zr phase in the dissimilar bonding.AcknowledgementsThis work was supported by the Department of Energy (grant number DE-FE0012272); theUniversity of Missouri System (award number FastTrack-16002R).783References[1] J.

D and Control Technologies for Computer-Aided Manufacturing of Laminated Engineering MaterialsSangeun Choi 1, Ravi Hebbar 2, Yong Zheng 3 and Wyatt S. Newman 4AbstractThis paper presents recent progress in software, material handling and tangent-cutting control in support of Computer-Aided Manufacturing of Laminated Engineering Materials(CAM-LEM). Progress in CAD focuses on the definition of a new layered file format for describing 3-D solids in terms of thick slabs with ruled-surface edges. For material handling, we present new algorithms for automatic generation of mask hole patterns used in selective-area vacuum gripping, which is required for our laminated assembly process. Finally, we present recent results of object fabrication using thick-slab, tangent-cut layers1.

Finally, we present recent results of object fabrication using thick-slab, tangent-cut layers1. Introduction Computer-Aided Manufacturing of Laminated Engineering Materials (CAM-LEM) is a sheet based approach to Solid Freeform Fabrication being developed at Case Western Reserve University [7, 8, 10, 11]. In this process, layers computed from cross sections of an object's CAD description are laser-cut from sheet material and subsequently stacked and laminated to assemble a part. Such stacks are post-processed to densify the material, fusing it into a monolithic solid. With this approach, we not only realize the original CAD description, but we also satisfy functional engineering properties (e.g. rigidity, strength, thermal expansion, surface finish, etc).One of the most important characteristics of a part fabricated from a successful SFF system is acceptable surface finish.

rigidity, strength, thermal expansion, surface finish, etc).One of the most important characteristics of a part fabricated from a successful SFF system is acceptable surface finish. With few exceptions [e.g., ShapemakerII [1], SDM [2], Stratoconception [3]], current SFF processes build parts from thin, vertically-extruded layers resulting in "staircasing" of the resulting surface. Commonly, one reduces the magnitude of the staircase roughness by reducing the thickness of each layer, but at the cost the fabrication speed.For our CAM-LEM system, a 5-axis laser cutting platform was built to translate and rotate sheet material under a stationary laser. By cutting contours at the appropriate tangent angles with this platform, it is possible to better approximate an object's surface and/or improve the build rate. However, our process of tangent-cutting and assembling layers presents multiple significant technical challenges. Progress in three of our challenge areas is presented here.

However, our process of tangent-cutting and assembling layers presents multiple significant technical challenges. Progress in three of our challenge areas is presented here. First, we describe how to represent the original CAD model in terms of an approximation using relatively thick slabs with ruled-surface edges. Second, we introduce improvements on our technique for extracting desired regions from the cutting table and precisely stacking them on the part assembly. Finally, we illustrate and discuss recent progress in building 3-D objects with our 5-axis CAM-:-LEM system.1 Department of Mechanical Engineering, Case Western Reserve University. 2 Department of Electrical Engineering, Case Western Reserve University. 3 Department of Electrical Engineering, Case Western Reserve University. 4Associate Professor, Department of Electrical Engineering and Applied Physics, Case Western ReserveUniversity, Cleveland, Ohio 44106 .6432.

4Associate Professor, Department of Electrical Engineering and Applied Physics, Case Western ReserveUniversity, Cleveland, Ohio 44106 .6432. Layered File Format (LFF) Representation of Solids To exploit the potential advantages of building objects with thick, tangent-cut layers, we must first derive an approximation of our original CAD model in terms of layers that can be fabricated with a practical system. Laser cutting is used in our system, but the same constraints apply to 5-axis CNC machining, 4-axis wire-EDM or hot-wire cutting and 4-axis water-jet cutting of sheet materials. For this geometric processing task, we preferably perform computations directly on the native CAD file, although our approach is applicable to STL format files as well.In our process, we distinguish "slices" from "layers." We define a "slice" as the 2-D cross section of our model at a given z-height.

A slice can be represented in terms of closed contours--"outer" contours enclosing regions of material, and "inner" contours describing punctures within the enclosed regions. Such contour information can be used in an alternative model representation, such as in the contour file formats SLC[4] and CLI[5]. A representation in terms of slices alone is adequate for systems that build with thin, vertical extrusions of layers. For thick, tangent-cut slabs, however, such representation is inadequate; it is also necessary to specify the shape of the edge surfaces of thick slabs. We thus define a "layer" as a 3-D body with planar upper and lower surfaces (slices) connected by edge surfaces. Objects can be defined in terms of stacks of such layers. Technically, such a layered representation can identically redescribe any CAD model.

Objects can be defined in terms of stacks of such layers. Technically, such a layered representation can identically redescribe any CAD model. However, the class of layers that can be fabricated by a line-of-sight cutting process is restricted to those with edge surfaces defined in terms of ruled surfaces [6]. Layer bodies with ruled-surface edges are simply described by one-to-one mappings of each contour on the upper surface to a corresponding contour on the lower surface. The mapping includes a functional specification of the path of a span-connecting upper and lower contours--swept along the perimeters. The swept path corresponds identically to the cutting-tool path for fabricating the layer, thus guaranteeing feasibility of fabrication.Since a ruled surface cannot identically match an arbitrary surface, our thick-layer representation with ruled edge surfaces can only approximate an exact CAD model.

Since a ruled-surface description joining two contours is not unique, the space of alternative ruled surfaces should be to the original model. An algorithm for performing this explored to identify the best fit computation based on an energy-relaxation technique was introduced in [7] and is further detailed in [8]. Having computed an optimal reconstruction that can be fabricated by a line-of-sight cutting process, the result should be stored in a format that is compact, precise, and efficient to interpret on-line during fabrication. For this purpose, we define a layered file format (LFF), as illustrated in Fig 1.In an LFF file, a part is represented by successive parallel layers in ascending order in the vertical direction. Each layer is defined by a bottom slice, top slice, and ruled layer edge-surfaces between the two slices. Slices are described by the boundary contours, which can be composed of polylines as well as higher-order curves.

Slices are described by the boundary contours, which can be composed of polylines as well as higher-order curves. An LFF file can be generated by slicing CAD models and stitching upper and lower cross-sectional contours with ruled surfaces, or by interpreting reverse engineering data collected by surface digitizing or from Computed Tomography (CT) imaging.Since our layers describe true 3-D bodies, we can also associate material descriptors with layer bodies. To represent mixed material composition within a layer, we define a layer in terms of a number of regions. Each region is a 3-D body described in terms of a closed (possibly punctured)644surface description, and each closed surface contains a labeled material. The regions can be adjacent to each other or disjointed.Fig. I illustrates the format.

.. I Contour k I .. , I.. , I Contourj ~tMaterial property,Material property, if being an outer·edge5, vs.ifbeingan -.... (surface)5, vs.. .Edge I (surface)Edge 2outer-edgeSbSb~+1 . ., I Contour m I . .. I Contour n I . " I+1 . ., I Contour m I . .. I Contour n I . " ~3 kinds of curves a) Polyline: described by vertices; b) Elliptic arc: described by cente r, major axis, minor axis, directio n, start-angle and end-angle; c) Spline: described by degree, co points, knots, and weights.Curve I Curve 2ntrolencode we Finally, contour descriptions in terms of "curves." Three types of curve entities have been defmed in the current version of the LFF file format. They are the the elliptic are, polyline, spline and curve, the Non-Uniform including B-Splines Rational all (NURBS).

They are the the elliptic are, polyline, spline and curve, the Non-Uniform including B-Splines Rational all (NURBS). contours in our description lie in parallel planes, only 2-D curve descriptors are needed.SinceFigure 1 : LFF representation of a solid modelA polyline is a series of connected straight line segments. A polyline with n segments is defined by its n+ I 2D vertices given in x-y coordinates.An elliptic arc is part of an ellipse. It is defined by a center point, a major axis vector, a length ratio of the minor axis to the major axis, a direction flag, a start-parameter, and an"end-parameter (analogous to start and end angles). A circular arc is a special case of elliptic arc when the length ratio equals 1.0. The elliptic arc is parameterized. Given the center C, major axis M and minor axis N, a point P on the ellipse can be located by parameter u as P(u) = C + M cos(u) + N sin(u).Spline curves, rational and non-rational, are also parameterized.

A spline is defined in terms of a specified degree, control points, knots, and (for rational splines) a set of weights. (see [9] for details).Our LFF file format achieves our stated objectives. It can represent contour descriptions within a slice plane precisely and compactly through the use of higher-order functions. It can also encode material information throughout the volume of a body. It represents tangent information for edge645surfaces in terms of a compact representation (scalar functions defining ruled surfaces) that encodes the full precision achievable by a line-of-sight cutting process. Finally, and most importantly, the format guarantees that each layer can be fabricated using a line-of-sight cutting means, and the specification decomposes efficiently for real-time interpretation for machine control. We consider this format to be ideal for sheet-based processes, including our CAM-LEM system.3. Selective-Area Material Handling of Layers.

We consider this format to be ideal for sheet-based processes, including our CAM-LEM system.3. Selective-Area Material Handling of Layers. A second technical challenge of the CAM-LEM approach is material handling. Since we cut each layer individually, the desired regions must be extracted from the cutting table and assembled precisely to form the desired part. For this task, we utilize a novel masked vacuum gripper In previous work, we demonstrated that it is feasible to extract technique (see [10] for details). selective regions using a vacuum gripper masked by a selectively-punctured, non-porous material. In recent progress, we have improved and automated the mask-hole pattern definition, we have extended the mask design to automated computation of multi-purpose masks, and we have proposed a means to extract tangent-cut regions from waste material.Our procedure for automated mask generation is described here.

Consistent with our LFF layer descriptions, we refer to the closed contours defined on the upper slice plane of each layer. (Only the upper slice plane will come in contact with the vacuum gripper). This slice consists of one or more outer contours, each possibly enclosing one or more inner contours. In our mask generation process, simply stated, each outer contour is successively eroded, each inner contour is successively dilated, and mask holes are defined at fixed intervals along the resulting contours. Ideally, this results in a sequence of closed, non-intersecting contours describing a uniform distribution of mask holes that can be punched along efficient trajectories.While conceptually simple, complications occur with this approach. For a non-convex contour, erosion eventually leads to self-intersecting contours, as illustrated in Fig 2. (A similar problem can occur with dilation of inner contours).

For a non-convex contour, erosion eventually leads to self-intersecting contours, as illustrated in Fig 2. (A similar problem can occur with dilation of inner contours). To resolve this problem, we analyze each eroded contour to test for self intersections. Self intersections define multiple closed loops. Each such closed loop is tested for its direction of rotation. If the direction of rotation is consistent with the direction of rotation of the original contour, then the resulting loop is retained. Reversed circulations indicate contours to be deleted. In a similar manner, the problem of intersections among multiple eroded/dilated contours is solved by locating the points of intersection of one contour with another and then retaining those inner/outer contours consistent in direction, as illustrated in Fig 3.

Incorporating these considerations, our algorithm is capable of generating correct mask-hole distributions for arbitrarily complex regions, resulting in fill patterns along concentric, closed contours. The corresponding mask production time is faster and more precise than a raster scan, analogous to the comparison between vector graphics and raster graphics. (This same algorithm may be applicable to efficient scan patterns for point-wise SFF building techniques as well.)646H\ORIGINAL CONTOURERODED CONTOURDISCARDED CONTOURS RESULTING FROM SELF INTERSECTIONRETAINEDMASK HOLE PATTERNFigure 20 { ) DORIGINAL CONTOUREROSION I DILAnON Figure 3RETAINED CONTOURSMASK HOLE PATTERNA limitation of our masked vacuum gripper technique is that production of masks can be too time consuming if a separate mask is required for each layer. In a recent extension, we automatically generate masks applicable to multiple layers.

In a recent extension, we automatically generate masks applicable to multiple layers. The motivation is obvious for gradually changing corresponding similar. "~~;; However, multi-use masks can also surprisingly be dissimilar layers. The procedure for generating multi-use masks is illustrated in Fig 4. The first two shaded regions are the desired areas The third image to be grasped. shows the Boolean intersection of the two desired regions. The mask pattern applicable for grasping both regions, the fourth image, is generated by our generic technique, but based on the contours of the intersection of regions.since 11{1¥~l':-~~ masks would 11I~lli ;;:~ 'I" \\'. 1111 ,,'. 11K' 11,11 ...J::: :::~l ~~.J=-%)lllayers,the alsobe11~111employedforSLICE 1SLICE 2 FINTERSECTION OF TARGET REGIONSMASK HOLE PATTERN4IgureA multiusable mask once generated should be checked for feasibility.

An analysis of the force exerted by the gripper on the part is done by solving for Newton's force balance and Euler's moment balance equations . The analysis requires the gravity force on the slice, the centroid, suction force per mask hole, the location of the mask holes and the number of mask holes, which are all known. The only unknown is the reaction pressure distribution between the grasped part and the mask. The space of the possible reaction pressure distributions is approximated by a finite number of hypothetical concentrated reaction forces distributed over the interior of the target region. In proposing reaction force coordinates, there is no penalty for including an excess number of locations. To assure a complete flooding of candidate reaction force locations the same mask generating algorithm is employed.

To assure a complete flooding of candidate reaction force locations the same mask generating algorithm is employed. The result is used as input to a linear programming equation solver, which solves for force/moment equilibrium feasibility.647FIVE SLICES OF A FLUIDIC DEVICE Figure 5INTERSECTION OF TARGET REGIONS Figure 6MASK HOLE PATTERNThe above method was employed to test the feasibility of a multiusable mask for five slices of a fluidic device, shown in Fig 5. The multiusable mask for the five slices, shown in figure 6, was fabricated and tested, and it successfully picked up each of the five layers while leaving the waste material behind in all cases.A third recent advance in our material-handling challenge is the extension of our masked-gripper technique to manipulating tangent-cut sections. Figure 7 illustrates the potential problem of extracting a desired section of material when tangential cutting is employed.

Figure 7 illustrates the potential problem of extracting a desired section of material when tangential cutting is employed. The figure represents a side view of a layer of a tilted cylinder cut from sheet stock, still surrounded by waste material. is not possible to It the desired section vertically extract without the waste material, nor is it possible to first extract the waste material vertically without interfering with the desired region. To solve this problem, we propose a second pass cutting trajectory in which the laser cuts the waste material as shown in the top view of Fig 7. The second-pass cut is oriented vertically, and it the upper contour of the desired region over those path segments bounding "overcut" regions (sections for which the surface normal of the layer's tangent-cut edge surface has a downward-facing component, thus blocking waste material from vertical extraction). The second pass cuts only through waste material.

The second pass cuts only through waste material. The result is that the waste material exposed to the top surface is then extractable vertically without interfering with the desired region. After the waste is extracted and discarded, the desired region is removable by vertical extraction without interference. The target slice can then be easily extracted, and what remains on the cutting table is discarded. This technique is still under development and testing, but it appears to be general enough to handle arbitrarily complex cases.WASTE MATERIAL TARGET SLICEI CUrING TABLE~,,~ICROSS SECTIONAL VIEW OF FIRST PASS CUTTINGSECOND. PASSC.UT.T.IN.~ . '.4 ~ ..........W2EXTRACTION..

PASSC.UT.T.IN.~ . '.4 ~ ..........W2EXTRACTION.. [Mil'Iinterfering withCROSS SECTIONAL VIEW OF SECOND PASSCUTTING AND WASTE EXTRACTIONTARGET EXTRACTION~RESIDUE ON THE CUTTING TABLECROSS SECTIONAL VIEW OF TARGET EXTRACTION AFTER WASTE EXTRACTIONFIRST PASS CUTTING TRAJECTORYSECOND PASS CUTTING TRAJECTORYBOTTOM CONTOURTOP CONTOURTOP CONTOURBOTTOM CONTOURSPANSfollowsTOP VIEW OF SECOND PASS CUTTINGTOP VIEW OF FIRST PASS CUTTINGFigure 76484. Tangent-cut Parts Fabrication Our CAM-LEM tangential cutting system is comprised of 3 parts: a cutting tool, a cutting platform and a material handling module. Our cutting tool is a 50W CO2 laser mounted vertically above the cutting table. The cutting platform consists of 3 translational axes (x,y,z) and two rotational axes (roll and pitch). This mobility enables the mechanism to position an arbitrary point on its cutting surface at an arbitrary orientation (within joint constraints) at the laser focal point.

This mobility enables the mechanism to position an arbitrary point on its cutting surface at an arbitrary orientation (within joint constraints) at the laser focal point. Our 5 axis mechanism carries a cutting table that clamps sheet feedstock to an aluminum honeycomb cutting surface by drawing a vacuum beneath the cutting surface. Material handling (loading, picking-up and stacking) is executed by 3-axis motion of a Cartesian robot with our selective area vacuum gripper design as its end-effector.Recent results of tangential cutting performance based on thick-slab, ruled-surface layered approximations of solid objects are presented here. Figure 8 shows the fabrication of a geometric test object consisting of the union of a 60mm-high cone with a 30mm-high rectangular prism. The object was defined in a CAD modeller, sliced at 11 slice planes, and redefined in terms of 10 layers with ruled-surface edges.

The object was defined in a CAD modeller, sliced at 11 slice planes, and redefined in terms of 10 layers with ruled-surface edges. The functional relationship defining the ruled-surface edges was used to define the path of the cutting tool. Motion of the cutting platform was derived from this tool path through inverse kinematics, as described in [11]. The ten layers were cut from 6mm (Our models were assembled by hand, since our automated material sheets of polystyrene foam. handling concept for tangent-cut layers has not yet been implemented in our system).Since each layer of the test object was defined in terms of ruled edges, a feasible tool path was guaranteed. Each ruled surface was approximated in terms of interpolation between more than 500 tabulated spans. Although this data constituted a feasible and accurate model, some trajectory modifications were required at run time.

Although this data constituted a feasible and accurate model, some trajectory modifications were required at run time. In particular, some of the spans near complicated regions required high joint velocities and accelerations to maintain a constant material removal rate. To the prescribed paths were avoid dynamically-undesirable or preprocessed before execution to identify problem conditions. Cutting speeds were scaled down linearly (preserving the same path shape) in such regions to satisfy dynamic constraints. A consequence of this time scaling is that it was not possible to maintain a constant material removal rate, yet the laser power remained fixed. As a result, some overcutting (excessive erosion) occurred in regions where the trajectory speed was reduced. This problem should be relieved by coordinating laser-power modulation with cutting speed.infeasible joint commands,A second test part, a 22-layer reconstruction of a head, is shown in Fig 9.

This problem should be relieved by coordinating laser-power modulation with cutting speed.infeasible joint commands,A second test part, a 22-layer reconstruction of a head, is shown in Fig 9. The tangent-cut layered head is compared to the corresponding 22-layer reconstruction using vertically-cut layers.The fabrication parameters for both of the tangent-cut test parts shown are summarized in Table 4.1. The present examples deliberately illustrate exaggerated cases of the advantages of tangent cutting. All 22 layers of the more complex test part (the in under 15 minutes using the quoted head) could be cut parameters (but In practice, thinner layers would be used to achieve greater precision, with corresponding increases in build time. Nonetheless, significant advantages in build speed can be expected relative to SFF systemsignoring material-handling time).Figure 8: Simple, 10-layer geometric test part649employing point-wise build techniques.

Summary and Conclusions This paper has presented our recent progress in CAM-LEM. We have emphasized the value of approximating solids in terms of relatively thick layers with ruled-surface edges, as such a description compactly captures the full model resolution achievable by line-of-sight type sheet cutters. At the same time, such a description guarantees the feasibility of fabrication of each layer using a line-of-sight cutter. This realization leads to a natural file format, which we have described as our "Layered File Format" (LFF).

This realization leads to a natural file format, which we have described as our "Layered File Format" (LFF). We argue that an LFF representation is optimal for sheet-cutting processes in terms of compactness, precision and ease of conversion to machine control.We have also described our advances in automated mask generation for our masked vacuum gripper material-handling technique." Our new algorithm for mask design results in concentric, closed contours, leading to a uniform distribution of mask holes as well as an efficient trajectory for punching mask holes. The gripper technique is further extended to multi-use masks. Experiments verified the value of the approach using significantly dissimilar layers.Finally, we have shown recent results from our 5-axis laser-cutting CAM-LEM system. Examples illustrate the value of tangent cutting, either in terms of surface fmish or in build speed.

Examples illustrate the value of tangent cutting, either in terms of surface fmish or in build speed. Although the examples shown were fabricated from styrofoam, the process is applicable to engineering650materials. objects from advanced ceramics and powdered metals.In ongoing work, we are developing procedures for fabricating tangent-cut, laminatedAcknowledgments This work was supported by the National Science Foundation under NSF grant DMI-94-20373, and by the Cleveland Advanced Manufacturing Program and industrial sponsors. This support is gratefully acknowledged. The work reported here is part of a multi-disciplinary program, and the authors recognize the contributions of team members professor James Cawley, professor Arthur Heuer, and Der-Lin Chow.References 1. Lee, C. H., T. M. Gaffney, and C. L. Thomas, "Soft Tooling for Low Production Manufacturing of Large Structures," in Solid Freeform Symposium Proceedings, University of Texas at Austin, Austin, Texas, pp.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA CAD-BASED WORKFLOW AND MECHANICAL CHARACTERIZATION FOR ADDITIVE MANUFACTURING OF TAILORED LATTICE STRUCTURESP. Koch*, H. Korn†, R. Kordass†, S. Holtzhausen*, C. Schoene*, B. Mueller†, and R. Stelzer*Institute of Machine Elements and Machine Design, TU Dresden, Dresden, Germany † Fraunhofer Institute for Machine Tools and Forming Technology IWU, Dresden, GermanyAbstractLattice structures are highly recommended for lightweight applications and cost reduction in additive manufacturing (AM). Currently, parts with lattice structures are still mainly used for illustrative purposes and rarely in industrial products. One important reason is that, due to their high dependency on macro- and micro-geometry, the mechanical properties of manufactured structures are difficult to predict.

One important reason is that, due to their high dependency on macro- and micro-geometry, the mechanical properties of manufactured structures are difficult to predict. Thus, even and precise struts are needed. In this paper, a workflow for fabrication of lattice structures with strut-diameters from 150 µm to 400 µm on commercial laser beam melting (LBM) systems is presented. Based on a CAD-integrated user-interface for lattice design, a customized slicing algorithm determines database-aided suitable exposure parameters which ensure that the properties of the manufactured struts will just be as specified upon design. Subsequently, compression tests are performed in order to verify the established workflow.

Subsequently, compression tests are performed in order to verify the established workflow. The developed tool enables designers to integrate AM-specific geometries into their components with little specific experience in AM.IntroductionAdditive manufacturing processes such as LBM offer significantly greater scope than conventional manufacturing processes with regard to the geometry of the product. A wide variety of shapes can be produced without the need to set up the machine or make any significant adjustments to the process.For solid components it is usually sufficient to use the standard parameters provided by the machine manufacturers. This applies both to the generation of slice data and to the production itself. Generally, within each layer the geometrical contours are determined and subsequently filled with hatches (contour-hatch exposure, CH).

Generally, within each layer the geometrical contours are determined and subsequently filled with hatches (contour-hatch exposure, CH). However, if the smallest part dimension approximates the laser focus diameter, the resulting geometry increasingly deviates from the nominal one when using the CH exposure [1,2]. This is particularly problematic with lattice structures, since their struts are usually very filigree. For example, own preliminary tests have shown a minimum achievable value of 400 µm for a targeted strut diameter of 200 µm when using CH-exposure [3].Since the effective mechanical properties of lattice structures disproportionally depend on their (heavily deviating) geometry, it is a rather futile task to tailor or even predict these properties for specific use cases. This is one of the main obstacles on the way to integrating lattice structures into industrial products [4].

This is one of the main obstacles on the way to integrating lattice structures into industrial products [4]. To solve this problem, uniform and precise struts that closely match the nominal geometry are needed.782This paper presents a workflow that covers the entire process chain from design to production of precise lattice structures even at strut diameters close to that of the laser focus. A custom CAD add-in for lattice design as well as a tailored slicer that uses point-like exposure and a database for exposure parameters are shown. Furthermore, compression tests are executed in order to verify the mechanical properties.Lattice DesignThere are several commercially available software solutions that feature lattice structures for additive manufacturing, most notably Netfabb by AutoDesk and 3-matic by Materialise. These are all stand-alone software products that require a part design to be frozen before lattice structures can be integrated.

These are all stand-alone software products that require a part design to be frozen before lattice structures can be integrated. Furthermore, they demand a significant amount of training and cause high license costs which render them uneconomical for small businesses. Meanwhile, even the smallest engineering companies own a CAD software that is used on a daily basis. Using this premise, we developed an add-in that enables designers to easily integrate lattice structures in parts directly within their familiar CAD environment and workflow.The first step is the definition of the volume which will subsequently be filled with lattice. Based on an existing volumetric body there are two ways of defining: creating a shell or extruding a sketch (Figure 1). The shell mode uses the user-defined wall thickness and open faces to determine the remaining solid body and (via a Boolean cut operation) the lattice volume.

The shell mode uses the user-defined wall thickness and open faces to determine the remaining solid body and (via a Boolean cut operation) the lattice volume. If the wall thickness equals zero, the whole body is filled and open faces are irrelevant. The sketch mode expects an arbitrary 2D sketch and two extrusion depths as input. Extruding the sketch in both normal directions according to the specified depths, a cutting body is created which is then used to determine the remaining solid and lattice volume via Boolean operations.With the volume now defined, the creation of the lattice structure itself is the next step. Therefore, the user selects a unit cell topology and size as well as the cell orientation. Given this input, the oriented bounding box of the lattice volume is calculated and meshed according to theFigure 1: Definition of lattice volume – top: via shell operation; bottom: via sketch extrusion783cell size.

The selected unit cell then is copied at each grid element that is at least partly inside the volume. Finally, struts exceeding the volume boundary are trimmed.As most CAD environments are based on Constructive Solid Geometry (CSG), which works by Boolean combinations of geometrical primitives, modeling the lattice will quickly crash the software due to the high number of strut intersections even in medium sized lattice structures [5–7]. To avoid this problem, the lattice structure will not be modeled, but the geometrical information is stored in a custom data structure in the backend instead. This data structure contains a list of nodes with their spatial coordinates and a list of struts connecting two nodes each. Additionally, each strut carries information about the geometry of its cross section.

Additionally, each strut carries information about the geometry of its cross section. To give the designer visual feedback on the just created lattice structure, a rough tessellation is computed and imported into the CAD environment (Figure 2).SlicingIn order to manufacture the designed parts on a LBM machine, slice data is needed. As described above, the standard slicing tools and exposure parameters provided by the machine manufacturers are usually optimized for massive parts and therefore not suitable for filigree lattices. Furthermore, these slicers expect triangle meshes of the parts as input which propagates the discretization error inherent in every triangulation to the slice data. When employing our lattice design approach the solid body is stored as CSG and the lattice as parametric data, both of which are exact representations.

When employing our lattice design approach the solid body is stored as CSG and the lattice as parametric data, both of which are exact representations. To make use of this major advantage we developed a custom slicer that consists of two modules: one for the solid body and one for the lattice structure.With the contour-hatch exposure being perfectly suitable for massive regions this strategy is adopted for the solid body. For this, the first step is the creation of a stack of parallel planes inside the CAD environment. The planes are oriented according to build direction and the distance is equal to the selected layer thickness. Intersecting the solid body with the plane stack the contour within each layer is determined as a sequence of straight lines, curves, and splines.

Intersecting the solid body with the plane stack the contour within each layer is determined as a sequence of straight lines, curves, and splines. This sequenceFigure 2: Lattice design via sketch mode – left: original body with selected sketch for extrusion (yellow); right: generated lattice visualized as triangle mesh784Figure 3: Slicing of a single BCC unit cell – left: slicing plane right: a) point exposure; b) collinear line; c) normal line; d) combined linesis then tessellated with straight lines, offset, and filled with hatches as needed. The tessellation will of course introduce a discretization error, but higher qualities can be achieved with significantly less memory overhead compared the common method of slicing triangle meshes.Slicing the lattice structure is not as straightforward, especially when targeting thin struts with diameters close to that of the laser focus.

Scanning the strut contour will already result in diameters diverging from the nominal value, as described above. One idea to avoid this deviation is to expose just a single point coinciding with the center of the strut cross-section in each layer and adjust the resulting strut diameter by varying laser power and speed [8]. We modified this idea: the exposure point is expanded to one (or two) short straight exposure line(s) that are either normal or collinear (or both) to the projected strut direction (Figure 3). The resulting strut diameter is then adjusted by varying the line length while keeping the laser parameters constant.

Instead, the needed line lengths are computed automatically using a parameter database of previously executed build jobs (Figure 4). Line lengths are always computed simultaneously for all struts with the same inclination to the building plane. Firstly, of all available parameter sets the ones that match the current powder material, layer thickness, and strut inclination are selected. Sets with bad strut quality are discarded at this step, too. Secondly, the remaining parameter sets are checked if there are any that match the laser parameters previously used for slicing the solid. If yes, these sets are selected for further processing. If not, all the remaining parameter sets are grouped by laser power, speed, and focus diameter. The group with the most data points is selected, with achieved strut quality serving as a tiebreaker. A quadratic function is then fitted to the data point, linking used line length to resulting strut diameter.

A quadratic function is then fitted to the data point, linking used line length to resulting strut diameter. With this relation it is possible to calculate a suitable line length for each strut of the current build job.The qualitative and quantitative advantages of Pseudo-P-exposure (left) over CH-exposure (right) can be seen in Figure 5. Both lattice structures are manufactured on the same LBM machine using the same parameters: Ti-6Al-4V, 25 µm layer thickness, 625 mm/s laser speed, 100 W laser power, and 150 µm laser focus diameter. For Pseudo-P the line length is 141 µm and for CH the targeted strut diameter is 200 µm. The Pseudo-P-exposure leads to noticeably finer and more even struts.Compression testsTo verify the described workflow for lattice structure manufacturing, test specimens are made and subjected to compression tests.

The determined effective mechanical parameters are compared to other test results available in-house or in corresponding literature [9].The specimen geometry is based on a basic body-centered cubic unit cell (bcc) with varying edge length L and strut diameter D (Figure 7 left). This unit cell is repeated 10x10x15 times to form a standing cuboid (Figure 7 right). The edge length L is either 1, 2, or 3 mm and the diameter D is ranging from around 180 µm to 800 µm as a result of varied line lengths. These786Figure 7: Compression test specimen – left: bcc unit cell; right: specimen geometry, consisting of 10x10x15 unit cellsvariations lead to relative densities from 2% to 22% of the test specimens compared to block material.An example of an executed compression test is shown in Figure 6. The test conditions are in accordance to German standard DIN 50134 and compression speed set to 0.1% per second (constant, based on the undeformed height of the specimen) [10]. With max.

The test conditions are in accordance to German standard DIN 50134 and compression speed set to 0.1% per second (constant, based on the undeformed height of the specimen) [10]. With max. test forces of <5 kN on a 50 kN test machine, machine frame deformation is considered negligible and therefore the displacement of the actuator is used to measure the strain. As the platens did not show dentations and gliding of the specimens is visible, uniform load distribution is assumed.Figure 6: Compression test of specimen with L = 3 mm and D = 250 µm left: initial state; right: compressed to block787Figure 8: Young’s modulus of specimens manufactured using Pseudo-P-exposure, CH-exposure, and results of corresponding testsIt is noticeable that even in the compressed state the cell structure is visible, which indicates a uniform compression due to homogeneous struts.

This is matching the findings related to strut quality described above.The measured Young’s moduli as a function of relative density are depicted in Figure 8. Added are two data points of similarly tested lattice structures that had been manufactured using CH-exposure and four data points taken from [9]. It is evident that our proposed method of exposure results in lattice structures that are in line with the mechanical properties of those manufactured using established strategies or by other researchers. The advantage of our method is the wider range of possible modulus values, especially towards very flexible and fragile structures as well as higher manufacturing speed due to simpler scan line geometry.ConclusionIn this paper we presented a workflow to design lattice structures for additive manufacturing directly within a CAD environment and integrate them into existing parts.

The structures can be created using two different modes of volume definition and with selectable size, topology, and orientation. Upon creation, the geometrical information of the lattice is stored in a data structure in the backend and a rough triangle mesh is generated for visualization purposes.After lattice design, a custom slicer is employed in order to calculate suitable slice data. First, the massive regions of the part are sliced and the scan vectors are determined using contour- hatch exposure. As this exposure strategy yields insufficient results when dealing with filigree lattice structures these are sliced using an exposure strategy consisting of short straight lines with788precisely defined length.

To calculate the required line length, a database of previous build jobs is employed and considering solid as well as strut related input suitable laser parameter sets are selected and the data points are fitted with a quadratic function.The executed compression tests show that using our exposure strategy results in lattice structures that are on par with other strategies regarding structural and mechanical properties. Furthermore, a wider range of properties is achievable, especially towards very flexible structures, but also when aiming at higher stiffness.Using the developed tools any designer can integrate lattice structures in arbitrary parts even with little AM-specific knowledge and only having prevalent CAD software at hand. In the future more features such as graded lattices will be implemented and the usability raised.

In the future more features such as graded lattices will be implemented and the usability raised. Additionally, a guideline is planned on how the parameter database can best be filled to yield sufficient results for other materials or LBM machines with as few as possible build jobs.AcknowledgementsThe research presented in this paper was funded by Deutsche Forschungsgemeinschaft(DFG, German Research Foundation) – 289647464.References[1] Dadbakhsh, Sasan, Speirs, Mathew, Kruth, Jean-Pierre, and van Humbeeck, Jan. “Influence of SLM on shape memory and compression behaviour of NiTi scaffolds.” CIRP Annals Vol. 64 No. 1 (2015): pp. 209–212. DOI 10.1016/j.cirp.2015.04.039.[2] Ghouse, Shaaz, Babu, Sarat, van Arkel, Richard J., Nai, Kenneth, Hooper, Paul A., and Jeffers, Jonathan R.T. “The influence of laser parameters and scanning strategies on the mechanical properties of a stochastic porous material.” Materials & Design Vol. 131 (2017): pp. 498–508.

This presentation will explore path planning for efficient imaging of printed geometries. The acquired scans are then polygonized and compared to the desired geometry which is utilized in defining machining areas with excess material or underbuilt areas for additional material deposition via gas metal arc welding (GMAW).IntroductionA current challenge with three-dimensional scanning of complex WAAM printed geometries is the visibility of every feature and surface of the geometry. Current static configurations of scanners inhibit full discernibility of an entire part. Current research is evaluating part positioner angles and positions that create the most ideal configurations for scan images. In this paper, a three-blade geometry is scanned and analyzed using a ZEISS GOM ATOS Q fringe projection scanner. The ATOS Q is mounted to an overhead fixture that can translate laterally and vertically for generating desired scan trajectories.

The ATOS Q is mounted to an overhead fixture that can translate laterally and vertically for generating desired scan trajectories. The part positioner is rotated and tilted for the scanning process to optimize the field of view of the scanner for part geometry. Generation of the requisite scan path and imaging sequence is non-trivial, however, as obscurations of the part geometry and surface features, particularly printed layer beads, introduce challenges leading to redundant scan areas and a prohibitively large number of scans. Quality scans are characteristically easier to produce via manual, handheld scanning but are substantially more labor and time intensive.Hybrid Cell ProcessUT’s hybrid cell begins by using a Kuka KR50 welding robot with a coupled part positioner to print a part geometry onto a pallet mounted to a KP-2 rotary-tilt part positioner. The part, pallet and tilt stage are translated on a KL4000 linear positioner to the metrology area of the workcell for scanning.

The part, pallet and tilt stage are translated on a KL4000 linear positioner to the metrology area of the workcell for scanning. The GOM ATOS Q scanner incrementally scans across the part surfaces, scans that are then stitched together in GOM Inspect to create a 3-dimensional point cloud. This point cloud can then be converted into a water-tight STL CAD model. The scanned geometry is then compared to a CAD model of the intended geometry and is used to inform the machining process of overbuilt758areas in need of machining, or underbuilt areas in need of additional deposition. The part and pallet can then be translated back to the KR-50 for additional printing or can be transported into a five- axis Haas UMC-750 machining center via a Kuka KR250 robot. Post-machining, the part can be returned to the metrology area to be re-scanned by the ATOS Q.

Post-machining, the part can be returned to the metrology area to be re-scanned by the ATOS Q. The secondary scan informs the cell of any additional deposition or machining required and the process can be repeated until the final comparison is within tolerances when compared to the desired geometry.Scanner Operation The GOM ATOS Q scanner works by fringe projecting narrow-band blue light across the surface of an intended target and the camera takes an image of the target with the overlayed pattern. The projector and lens of the camera are at a known angle to each other which allows for easy triangulation of the pixels of pattern images taken. The lens magnifies the area being measured and condenses pixels for finer detail scanning. The measuring area of the ATOS Q used in this cell has a working distance of 490 mm from the camera to the part surface. The sensor captures up to 2 x 12 million coordinate points per scan.

The measuring area of the ATOS Q used in this cell has a working distance of 490 mm from the camera to the part surface. The sensor captures up to 2 x 12 million coordinate points per scan. The scan volume for this scanner is variable between 100mm x 70mm – 500mm x 370mm. After a projection occurs, the camera captures an image of the patterned surface, and that image is converted into a STL format with the help of reference points. When multiple scans are conducted, the respective images are stitched together. Areas of overlap are verified against one another, and new areas are added on. At the end of the scanning process all the images are “polygonised”, which means they are all combined into a single STL.Scanner ConfigurationAs displayed in Figure 1, the scanner is located in a fixed position above the linear track and part positioner. Once the part positioner moves the part into view of the blue light scanner, the scanner is instructed to take an initial scan.

Once the part positioner moves the part into view of the blue light scanner, the scanner is instructed to take an initial scan. The scanner determines its subject’s position via reference points on the pallet and displays its respective field of view in the accompanying GOM Inspect software. With the scanner aimed directly downwards, legible scans of the top of the part are easily attained as shown in Figure 1. This is due to the ability of the blue light to evenly distribute and scan directly across the surface with minimal shadows, feature obstructions and surface slopes.759Figure 1: Static Configuration of Scanner with Respective ScanThe underside of the blades, however, presents challenges as the part positioner must be tilted and rotated to appropriately be in view of the scanner (Figure 3). This causes light illumination to then strike the part at a shallow angle instead of directly perpendicular to the surface.

This causes light illumination to then strike the part at a shallow angle instead of directly perpendicular to the surface. The blue light striking at an angle leads to data dropout caused by shadows created by irregularities in the geometry, the base plate, or the light missing the part altogether. The underside of the blade requires a near 90-degree tilt of the part positioner. However, when the part positioner is tilted beyond an 85 degrees the base plate begins to overshadow the part from view of the scanner. The scan of the underside of the three-blade geometry is displayed in Figure 3. The areas of blue indicated data dropout, or areas where the certainty of the data collected cannot be verified. The resulting part geometry required 535 scans, which causes the scanning process to be longer than printing or machining. Printing the three-blade part requires approximately 3-4hrs, while machining requires 2hrs.

Printing the three-blade part requires approximately 3-4hrs, while machining requires 2hrs. Manual part scanning, moving the part positioner incrementally via the control pad, requires 6hrs, nearly double the print time, and three times the machine time. Each scan requires the part to be moved into desired position, the scanner must be actuated, then the blue light scans across the surface and an image is created. Even once this process is automated, each scan takes roughly 30 seconds to complete. This also does not include the time required to polygonise the scans together into one CAD model. Note that the time necessary for polygonising increases with the number of scans administered.760Figure 2: Scanning of Underside Of Three Blade Geometry.To combat this, various configurations of the scanner were examined for ideal imaging.

These studies determined that when the camera was tilted at a -25-degree angle, the underside scan of the part dramatically improved without significantly complicating or reducing the quality of the top surface scan of the blades. This configuration presented similar data drop out, but dramatically reduced scan time. Figure 4 displays the result of less than 200 scans in this configuration. A culmination of 200 scans requires a little over an hour and a half to complete at the rate of 30 seconds per scan. This reduces the amount of time required for adequate scanning by over a factor of 2.Figure 3: Underside Of Geometry Scanned From -25 Degree Position Of Scanner.761Data AnalysisAfter printing, prior to machining the geometry presents different challenges than a machined part. Prior to machining, the layers of printing on the part cause for some anterior areas to be obstructed, leading to data dropout.

Prior to machining, the layers of printing on the part cause for some anterior areas to be obstructed, leading to data dropout. Fortunately, these layers are still able to be analyzed via the GOM software to determine layer height, thickness, etc. as seen in Figure 4. This data is useful to determining errors made during the printing process such as overheating, excessive cooling, or misalignment. Unit vectors are placed along a single “slice” of the scan to determine the layer height and thickness.Figure 4: Layer Height StudyParts post machining face their own challenges when scanning. Machined parts are difficult to scan due to the reflective nature of the material. When the light reflects from the part, the scanner is unable to analyze it accurately. The use of scanning spray greatly improves scan quality but is an improbable option for our cell given the objective of automation.

The use of scanning spray greatly improves scan quality but is an improbable option for our cell given the objective of automation. A study was conducted to aid in understanding of the accuracy of the scanner with varying surface finishes and configurations. Spheres of identical diameter (1 inch) but different finishes were scanned from the same configurations to analyze the effects of surface finish on scan quality. These sphere scans were then compared to a CAD model of a standard 1-inch diameter sphere (Figure 5). This study also provides a visual representation of the field of view of a specified configuration of the scanner in relation to a geometry.762Figure 5: Matte Finish Sphere Study Deviation Of Scan To Expected CAD In Mm.For each study, reference points were evenly plotted along a section of the sphere and used to analyze the deviation from the expected data and the data provided by the scan.

These reference points were exported into an excel file to be analyzed independently of the GOM software. Plots were created for each experiment comparing error of the scan data to the position along the sphere. This data provides some insight as to where data dropout is most likely to occur when scanning from each configuration.Figure 6: Error Vs. Location On Matte Finish SphereThe plot above (Figure 6) is an example of many derived from the data collected from each of the spheres. Notice the steeper the angle gets on the sphere the more error it experiences. This is to be expected as the sphere isn’t reflecting as much light back into the scanner, much of it is missing the sphere entirely.763Conclusions/SummaryIntegrating metrology into the hybrid manufacturing cell via fringe projection scanning greatly improves efficiency and quality of the product created. With an informed process via the 3D scanner the needs of a product are quickly identified and met.

With an informed process via the 3D scanner the needs of a product are quickly identified and met. The sphere study conducted concluded the exact field of view of the scanner in relation to a part and aids in determining an ideal scanning information that can be used for optimal path planning of the part positioner.Future WorkA current area of research being explored is the potential to use a Kuka KR250 robot to articulate the GOM ATOS Q scanner as a tool and path plan the most ideal scanning configurations for a specified geometry. This would allow for ease of access to complex geometries that would be nearly impossible to scan from a static configuration. Figure 7 below displays a simulation created via Octopuz software of the scanner mounted at a 90-degree angle relative to the gripper axis(?). Mounting at 90 degrees allows for more intricate scan positions without nearing joint limits of the robot. The shaded red box displays the field of view of the scanner.

Mounting at 90 degrees allows for more intricate scan positions without nearing joint limits of the robot. The shaded red box displays the field of view of the scanner. Notice the geometry on the part positioner is more complex than the previous three-blade geometry, but much of the geometry remains within the field of view of the scanner. The robot would also be able to move the scanner into an angle/position the combination part positioner and static configuration could not. Coupling the part positioner to the KR250 for scanner the same way it is conducted for welding increases the available degrees of freedom from three to eight degrees of freedom, which vastly expands our potential area of operation. In our hybrid cell, once the part finished machining it would be translated forward for scanning just as before. The KR250 would scan the part via the scanner, then place the scanner into a storage mount to the side.

The KR250 would scan the part via the scanner, then place the scanner into a storage mount to the side. The KR250 then picks the pallet up as normal and transports it for machining. Post machining, the pallet is returned to the scanning position via the KR250. The KR250 then releases the pallet, picks the scanner back up from the storage mount, and administers another round of scanning.Figure 7: KUKA KR250 with GOM Mounted ATOS Q Scanner764KeywordsHybrid manufacturing, wire-arc additive manufacturing, machining, fringe projection scanningAcknowledgments This work relates to Department of Navy award (Award number N00014-20-1-2836 in collaboration with Oregon State University) issued by the Office of Naval Research. The United States Government has a royalty-free license throughout the world in all copyrightable material contained herein.

The United States Government has a royalty-free license throughout the world in all copyrightable material contained herein. Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the view of the Office of Naval Research.765.

librating Large-area Mask Projection Stereolithography for Its Accuracy and Resolution Improvements Chi Zhou, Yong Chen* Epstein Department of Industrial and Systems Engineering University of Southern California, Los Angeles, CA 90089Corresponding author: yongchen@usc.edu, (213) 740-7829ABSTRACTSolid freeform fabrication (SFF) processes based on mask image projection such as digital micro-mirror devices (DMD) have the potential to be fast and inexpensive. More and more research and commercial systems have been developed based on such digital devices. However, a digital light processing (DLP) projector based on DMD has limited resolution and certain image blurring. In order to use a DLP projector in the large-area mask projection stereolithography, it is critical to plan mask images in order to achieve high accuracy and resolution.

In order to use a DLP projector in the large-area mask projection stereolithography, it is critical to plan mask images in order to achieve high accuracy and resolution. Based on our previous work on optimized pixel blending, we present a calibration method for capturing the non-uniformity of a projection image by a low cost off-the-shelf DLP projector. Our method is based on two calibration systems, a geometric calibration system that can calibrate the position, shape, size, and orientation of a pixel and an energy calibration system that can calibrate the light intensity of a pixel. Based on both results, the light intensity at various grayscale levels can be approximated for each pixel. Developing a library of such approximation functions is critical for the optimized pixel blending to generate a better mask image plan. Experimental results verify our calibration results.1.

Developing a library of such approximation functions is critical for the optimized pixel blending to generate a better mask image plan. Experimental results verify our calibration results.1. INTRODUCTIONCurrent market trends such as consumer demand for variety, shorter product life cycles, and higher product quality have resulted in the need for more efficient and robust manufacturing paradigm. Solid freeform fabrication (SFF) can fabricate parts directly from computer-aided design (CAD) models without part-specific tooling and fixtures. Thus it can significantly shorten product development cycle while satisfying customized design requirements.1.1. SFF Processes based on DMDThe large-area mask projection stereolithography is a SFF process based on mask image projection on a big surface area (e.g. 250x250mm). In this process, the 3-Dimentional (3D) CAD model of an object is firstly sliced by a set of horizontal planes.

250x250mm). In this process, the 3-Dimentional (3D) CAD model of an object is firstly sliced by a set of horizontal planes. Each slice is then converted into a 2- Dimentional (2D) image which serves as a mask corresponding to the layer to be built. The mask image is then projected by a light projection device onto a surface to form the layer of the object. By repeating the process, 3D objects can be formed on a layer-by-layer basis. This process can form a whole layer of multiple objects simultaneously and dynamically. Therefore, it provides a potentially faster approach that is independent from part size and geometry. It also removes the requirement of an accurate XY motion control subsystem. Therefore the process can be relatively inexpensive.Several research systems have been developed before based on the mask image projection process [1-12].

Therefore the process can be relatively inexpensive.Several research systems have been developed before based on the mask image projection process [1-12]. Both digital micromirror devices (DMD) and light crystal display (LCD) have been used as light projection device, and visible-light-cured photopolymer or UV curable resin have been used as the raw material. These research systems mainly focus on fabricating micro parts (hence mask projection micro stereolithography). In addition, commercial systems based on DMD have also been developed such as the Perfactory system from Envisiontec Gmbh [13] and the V-Flash system from 3D Systems Inc. [14].1.2. Pixel Blending for DMD-based SFF ProcessesThe mask image used in a mask image projection system is one of the most important process parameters. A well planned mask image is especially important for the large-area mask projection stereolithography.

A well planned mask image is especially important for the large-area mask projection stereolithography. Since a big platform area is required while the resolution of a DMD device is limited by the number of pixels (e.g. 1024x768), the pixel size of a projected image will increase accordingly. Therefore, the achievable accuracy and resolution will decrease in the built models.To address such a problem, an optimization pixel blending method [15] has been proposed. The method is based on the observations that: (1) the light beam of a pixel will spread to its neighboring pixels; (2) the light intensity of a pixel more follows Gaussian distribution; (3) the light intensity at any position is actually the sum of the light intensities contributed by all the neighboring pixels (defined as pixel blending); (4) a desired energy input can be achieved by intelligently manipulating the grayscale values of each pixel; (5) such exposure energy will then lead to a cured layer with desired shape.

We mathematically formulate the pixel blending problem in an optimization model, and use optimization tools to solve the problem. Both simulation and physical experiments have been carried out to validate the effectiveness and efficiency of the optimized pixel blending method. Figure 1 shows an example of dimensional accuracy study. Notice the traditional method will lose the small portion while the pixel blending method can get the shape quite close to the original one. Figure 2 shows the case of surface quality study. It is obvious that the building result based on the optimized pixel blending method has a much better surface quality.Figure 1. A simple test case for dimensional accuracy study [15].Figure 2. A simple test case for surface quality study [15].1.3.

A simple test case for dimensional accuracy study [15].Figure 2. A simple test case for surface quality study [15].1.3. Projection Image Calibration for Optimized Pixel BlendingAs demonstrated in our previous work, the pixel blending of light intensity would give us tremendous capability in selectively solidifying resin into a desired shape. However, in the pixel blending method presented in [15], we used one Gaussian function (at the center of the image) to approximate the light intensity of all pixels. This may not be suitable for a low cost off-the-shelf DLP projector. Due to the low cost optic components used in such a projector, the projection image may have distortion and the light intensity is non-uniform. These geometric and energy differences would affect our pixel blending results. In this paper, we present a method to calibrate such non-uniformity of a projection image.

These geometric and energy differences would affect our pixel blending results. In this paper, we present a method to calibrate such non-uniformity of a projection image. By incorporating the calibration results in our mask image planning framework, the part quality for the large-area mask projection stereolithography will be significantly improved.2. PROJECTION IMAGE CALIBRATION PRINCIPLEIn our large-area mask projection stereolithography system, we used an off-the-shelf DLP projector (CASIO XJ-S36) with 1024x768 pixels. The pixel size is around 0.265 mm when a platform is set at 271x204 mm. Considering its small size, it is rather difficult to use light measurement devices to directly measure the properties of a pixel.

Considering its small size, it is rather difficult to use light measurement devices to directly measure the properties of a pixel. In this paper, we propose a novel calibration method based on dividing the pixel properties into two categories:(1) Relative properties: For a pixel’s geometric properties such as its position, shape, size, and orientation, it is more important to measure the relative light intensity. That is, we can manipulate the incident light beam at our wish. As long as the measured light intensity after the manipulation is proportional to the incident light intensity, the measurement of such relative properties will be accurate.(2) Absolute properties: For a pixel’s energy properties such as its light intensity, no manipulation will be allowed. We have to measure the incident light beam directly.Accordingly, we develop two types of calibration systems: geometric and energy calibration systems. Different measurement methods have been developed in the systems.

Different measurement methods have been developed in the systems. Finally the light distribution of a pixel can be characterized by combining the geometric and energy calibration results.(1) Geometric Calibration Principle:We tried the method of mounting a camera directly toward the projected light beam of a DLP projector. However, the geometric properties of a pixel are hard to be measured form the captured images since it is difficult to remove the highlight at the surface of the lenses. To overcome the drawback, we develop a method based on measuring the light transmission from a translucent medium. That is, we add some translucent medium at the focus plane of a projected image. Hence, the incident light beam ( aK ) and transmission . From these optical equations, we ; and i tK ).

That is, we add some translucent medium at the focus plane of a projected image. Hence, the incident light beam ( aK ) and transmission . From these optical equations, we ; and i tK ). In addition, ( K can calibrate either portion of light intensity to compute the geometric properties of a pixel since they are proportional toiK ) is decomposed into three portions: reflection ( ; K KrK ), absorption ( +;KK∝K+K∝KK∝K=araiirittiK .rK . That is, the light shoots on the medium and a picture is taken by using the light reflected from the medium. However, in this method, a camera can not be mounted overhead the image; otherwise, it would block the light. The angle between the lens and the projection plane would result in large distortion in the picture.

The angle between the lens and the projection plane would result in large distortion in the picture. Extra efforts are required to correct such distortions.We tested the method of measuringtK , a camera can be directly mounted under the medium so no distortion exists in the pictures taken by the camera. The light transmitted from the medium ( tK ) can be measured directly. We use a digital microscope device (x50) to take a magnified picture for each pixel. The images can then be analyzed since the light intensity is approximately proportional to the grayscale level of the image. Accordingly the geometric properties of each pixel can be calibrated.Instead, in the proposed method based on measuring(2) Energy Calibration Principle:We use a photosensor and RC circuit to directly measure the light intensity of incident light. Since a photosensor has limited sensitivity and relatively big size (>3 mm), we can only measure the light intensity of a small region (e.g. 20x20 pixels).

Since a photosensor has limited sensitivity and relatively big size (>3 mm), we can only measure the light intensity of a small region (e.g. 20x20 pixels). Hence the light intensity of a pixel is computed by averaging the measuring result of the region that the pixel belongs to. Our method is based on the assumption that the light intensity changes only slightly within a small region (e.g. 2x2mm) while it may vary greatly from regions to regions. Based on our experimental results and the optical error analysis [16], we believe such an assumption is valid.By combining the geometric and energy calibration results, the light distribution of each pixelcan be characterized. An illustration of the whole calibration system is shown in Figure 3.Original colorful imageGray scale imageApproximated contourApproximated light intensity distributionImage ConvertorDirect Ellipse FitterPhoto SensorC a l i b r a t i o n r e s u l tCalibration SystemFigure 3.

Pixel calibration framework for a DLP projector.The remainder of the paper is organized as follows. In section 3, we present the geometric calibration system. In section 4, we describe the energy calibration system. The calibration results based on the presented systems are discussed in section 5. In section 6, we present the experimental results for verifying our result. Finally, we draw conclusion and discuss future work in section 7.3. GEOMETRIC CALIBRATION SYSTEM3.1. Calibration SetupThe setup of our geometric calibration system is shown in Figure 4. We use ProScope HR [17] with a 50 times lens to capture the image of a pixel. A high resolution XY linear stage is used to transform the microscope to cover the whole area of a platform. As mentioned before, some medium (paper, plastic, etc) is added between the microscope and the projector to avoid the highlight.

As mentioned before, some medium (paper, plastic, etc) is added between the microscope and the projector to avoid the highlight. Themicroscope is then placed directly under the medium such that the focus plane of the microscope is coplanar to the projection image.MediumMicroscopeProjectorLinear Motion SystemJoy StickFigure 4. Setup of our geometric calibration system.As shown in Figure 5, we developed a software system to integrate the linear motion of the microscope and its functions such as picture taking and saving. The client window can dynamically show the captured pictures. Initially the microscope is set at the original position of the projection image. At each position, we control the projector to turn on a related pixel and control the microscope to take a picture. Then image analysis algorithms are run in parallel to compute the geometric parameters. After that, the motion control system will translate the microscope to a new pixel position.

Then image analysis algorithms are run in parallel to compute the geometric parameters. After that, the motion control system will translate the microscope to a new pixel position. The procedure is repeated until all the pixels are measured and analyzed.Figure 5. Motion control and image analysis interface.3.2. Medium SelectionAs discussed in section 2, some media should be added for measuring light transmission. Thus, a related issue is to choose a proper medium to achieve the best measuring quality. Different kinds ofaK ) and transmission medium have been tried for a good balance between reflection ( tK ). After multiple tests, we select a white paper as the medium in our system. The white paper can ( give us good picture quality such as high contrast, good penetration, and low noise. Some other test results are also shown in Table 1. The black paper and blue paper absorbed too much light and the plastics reflect too much light.

Some other test results are also shown in Table 1. The black paper and blue paper absorbed too much light and the plastics reflect too much light. Thus they are not good candidates for measuring tK . Other translucent media is still possible; however, opaque or transparent ones are not appropriate.rK ), absorption (Table 1: Pictures under different mediums.White paperBlack paperBlue paperWhite plasticBlack plasticBlue plastic3.3. Image AnalysisThe pixel grayscales values in a digital image carry the intensity information. The grayscale value of a pixel varies from black (at the lowest intensity) to white (at the highest intensity). The pictures take by the microscope are represented in RGB colors (i.e. Red, Green, and Blue). To convert a color picture into a grayscale image, we can calculate the effective brightness or luminance of a RGB value. Based on it, we can then compute a grayscale value to match the brightness.

To convert a color picture into a grayscale image, we can calculate the effective brightness or luminance of a RGB value. Based on it, we can then compute a grayscale value to match the brightness. The approach we used is based on a commonly used conversion equation:grayscale = 0.299*r+0.587*g+0.114*b,where r, g, b is the red, green and blue components of the color image.As shown in the picture of “white paper” in Table 1, the environmental light has effects on the grayscale value of a pixel. That is, the background of the paper has the brown grain textures due to the environmental light. Our method to remove such effects is to take two pictures for each pixel exposure. That is, when moving the microscope to a position, we take a picture of the pixel exposure; after turning off the pixel, we then take another picture without the pixel exposure. The subtraction between the two pictures turns out to be the picture of the pure pixel exposure.

The subtraction between the two pictures turns out to be the picture of the pure pixel exposure. To reduce the noise, we may also take the two pictures for multiple times and use the average of all the subtraction picture as the final result. The above process is illustrated in Table 2. An alternative way is to take pictures in a darkroom to physically avoid the effect of the environmental light. Some test results are shown in Table 3.a.

An alternative way is to take pictures in a darkroom to physically avoid the effect of the environmental light. Some test results are shown in Table 3.a. The results are better than those based on the previous method.Table 2: Grayscale conversion and picture subtraction.A Pixel ExposureBackgroundSubtraction between “A Pixel Exposure” and “Background”e r u t c i pl u f r o l o Ce r u t c i P e l a c s y a r GTable 3: Grayscale conversion and picture subtraction.(a) Original colorful picture(b) Grayscale picture after low- pass filter(c) Above threshold(d) Boundary(e) Fitted ellipses(f) Ruler under microscope (in the same scale)1mmTo further remove the noise, we can also apply the low-pass filter to the grayscale image. The smoothed picture is shown in Table 3.b. We set a threshold 0.1 (the grayscale value is normalized between 0 and 1) as the cut-off value.

The smoothed picture is shown in Table 3.b. We set a threshold 0.1 (the grayscale value is normalized between 0 and 1) as the cut-off value. We then use a recursive algorithm to get the effective light region and the corresponding boundary which is shown in Table 3.c-d. Based on the assumption of Gaussian distribution, the contour for a Gaussian function is an ellipse. We use a fast non-iterative ellipse fit method, direct ellipse fit [18], to approximate the pixel shape. The parameters can then be computed (refer to Appendix A for the detail procedure). An example of the approximated ellipse is shown in Table 3.e. For a reference of the actual size, we also show an image of a ruler in Table 3.f. As shown in the figures, the real size of a pixel is about 1 mm, which is about 3~4 times of the ideal pixel size. This result validated the fact that the light beam spreads to the neighborhood pixels.3.4.

This result validated the fact that the light beam spreads to the neighborhood pixels.3.4. Geometric Calibration of the Whole PlatformWe apply the previous method to all the pixels in a projection image related to the platform. In our setup, the projection image has a resolution of 1024x768. Hence the above process needs to be run automatically in order to calculate the shape, size, orientation and position of each individual pixel. However, due to the intrinsic distortion, we found the project plane can not be adjusted into a perfect rectangle. The best one we use in our system is shown in Figure 6. Since our microscope has a limited image size (around 6 mm). We need to compensate the distortion in the motion control in order to capture the exposure of all the pixels.Figure 6. Dimension and shape of the image.4. ENERGY CALIBRATION SYSTEMThe geometric calibration system can capture the size, shape, orientation and position of a pixel.

Dimension and shape of the image.4. ENERGY CALIBRATION SYSTEMThe geometric calibration system can capture the size, shape, orientation and position of a pixel. However, it is hard to know the ratio between reflection ( aK ) and transmission iK ). Hence, we use light measurement sensors to directly measure the tK ) in the incident light beam ( ( light intensity of an incident light beam. A photoresistor can exhibit a resistance change in proportional to the exposed light intensity. Hence it can serve our purpose. An important consideration for selecting a photoresistor is the linearity of the output signal. An output that is linearly proportional to the input intensity (illumination) is highly desirable. We selected a photoresistor (350-00009 by Parallax [19]).

An output that is linearly proportional to the input intensity (illumination) is highly desirable. We selected a photoresistor (350-00009 by Parallax [19]). According to its technical specifications, the following equation shows the relationship between the resistance of a photoresistor and the illumination E:rK ), absorption (R AE a− ,=where R is the resistance of the device, A and α (sensitivity) are constants dependent on manufacturing process and material type. The photoresistor we used is made from type φ material and its. Based on the equation, we can easily compute the light intensity by measuring the0.8α= photoresistor’s resistance.The setup of our energy calibration system is shown in Figure 7. We use a microcontroller from Parallax [19] to integrate the measurement results of the resistance into the whole automated calibration system.

We use a microcontroller from Parallax [19] to integrate the measurement results of the resistance into the whole automated calibration system. We construct a simple RC circuit and use a RCTIME function provided by the microcontroller to measure the resistance of the photoresistor. To reduce the effect of fluctuation, we measured the resistance 100 times and take the average in computing the light intensity of each pixel.Photo SensorI/O PinCvssFigure 7. Setup of our energy calibration system.The measured light intensity of a pixel can then be used in computing the height A of itsGaussian function. That is, suppose the Gaussian function is:⎫ ⎤ ⎪ ⎥ ⎬ ⎥ ⎪ ⎦ ⎭ xσ and 0y ) is the center point of the pixel,⎧ ⎪ ⎨ ⎪ ⎩⎡ ⎢ ⎢ ⎣22())x − 0 2 σ xx(y − 0 2 σ yy1 2,−expA=)( , g x y+where ( 0x , yσ are used to characterize the shape of the 0y , 3 xσ , 3 yσ from the previous geometric calibration system. pixel.

CALIBRATION RESULTSThe calibration results based on the aforementioned calibration systems are presented asfollows. 5.1. Geometric Calibration ResultsOur initial experimental results showed that a pixel has a similar distortion with the ones in its immediate neighborhood. Hence, instead of measuring each pixel which will take quite a long time, we can only sample a portion of the pixels that are uniformly distributed. The properties of other pixels can then be interpolated from the calibrated pixels.We present the geometric calibration result of 10x10 = 100 sampled pixels for a grayscale value of 255 (full light intensity) as shown in Figure 8. For the sake of observation convenience, the size of the ellipses is magnified 300 times. The center point and orientation of each ellipse represent the center point and the orientation of the sampled pixel. We also perform the same experiments for lower light intensities.

The center point and orientation of each ellipse represent the center point and the orientation of the sampled pixel. We also perform the same experiments for lower light intensities. The results for the projection images with grayscale=200 and grayscale=150 are shown in Figure 9 and Figure 10 respectively.Figure 8. Distribution of 1 pixel (grayscale=255).Figure 9. Distribution of 1 pixel (grayscale=200).Figure 10. Distribution of 1 pixel (grayscale=150).Based on the results, we can draw the following conclusions and observation:(1) The pixels are non-uniformly distributed in the screen; (2) The distribution is consistent for different grayscale levels; (3) The shapes and orientations of the pixels are change gradually and smoothly in both vertical and horizontal directions, thus we can interpolate the properties of a pixel base on the calibrated ones in its neighbors;(4) The center portion has less distortion.

Thus using center portion of the image to build part canachieve better quality;(5) Pixel grid has distortion. The pixels are not located in the same line in both vertical and horizontaldirections and the whole screen is not a perfect rectangle.(6) The pixels are not evenly distributed in the projection image (with some shifting).5.2. Energy Calibration ResultsFrom the readout of the photosensor, we can compute the light intensities according to the resistor-intensity equation. A visualization of an experimental result on the pixels’ light intensity is shown in Figure 11. In the figure, different colors are used to represent the light intensity values. The light intensity is higher if the related pixel has a redder color; and the light intensity is lower if the pixel has a bluer color.(a) grayscale=255(b) grayscale=200 Figure 11.

The light intensity is higher if the related pixel has a redder color; and the light intensity is lower if the pixel has a bluer color.(a) grayscale=255(b) grayscale=200 Figure 11. Light intensity distributions for different grayscale levels.(c) grayscale=150Based on the results, we can draw the following conclusions: (1) The energy is non-uniformly distributed in the projection image; (2) The center of the image has higher light intensity; (3) The energy distribution is consistent for different grayscale levels.5.3.

Pixel calibration (grayscale=150).⎫ ⎪ ⎬ ⎪ ⎭6. VERIFICATION OF CALIBRATION RESULTS6.1 Shape VerificationTo verify the idea of pixel blending as discussed in Section 1.2 (i.e. the convolution of light intensity of neighboring pixels), we investigate 2x2 pixels for the full grayscale case (i.e. grayscale=255). We sample 5x5 different positions; for each position, we repeat the procedure as presented in the geometric calibration. The measured distribution results are shown in Figure 15.(a). We then compute the convolution of the 2x2 pixels based on the calibration result of a single pixel. The simulated result are shown in Figure 15.(b). A comparison of the two figures illustrates that the measured and simulated shapes are close. Hence it verifies the pixel blending idea.(a) Measured distribution for 2x2 pixels(b) Simulated distribution for 2x2 pixels.Figure 15.

Hence it verifies the pixel blending idea.(a) Measured distribution for 2x2 pixels(b) Simulated distribution for 2x2 pixels.Figure 15. Pixel blending verification.We also investigate 2x2 pixels for lower grayscale cases to show the consistence of the distribution. The results are given in Figure 16. As can be seen from the figures, the shape and the orientation are consistent with the single pixel case (refer to Figure 9 and Figure 10). Only the sizes are increased.(a) grayscale=200(b) grayscale=150Figure 16. Distributions of 2x2 pixels.We further performed physical experiments to verify the calibrated shapes. We spread a thin layer of resin on a red plastic plate, and then turn on 8x8 pixels at the left top corner and the left bottom corner respectively. The resin is solidified after several minutes. We clean up the liquid resin,mark the cured part with blue color, and then put them under a microscope. The magnified pictures are shown in Figure 17.

The resin is solidified after several minutes. We clean up the liquid resin,mark the cured part with blue color, and then put them under a microscope. The magnified pictures are shown in Figure 17. Based on the figures, we can clearly see the different orientations at different places. Further, the orientations are consistent with our calibration result as shown in Figure 8.(a) pixel at top left corner(b) pixel at bottom left cornerFigure 17. Physical experiment results for shape verification.6.2. Energy verificationSimilar physical experiment has been performed to verify the calibration result of the energy distribution. We turned on 4x4 pixels at the center and near the boundary (middle-right). We then expose them for the same period of time. The result is shown in Figure 18. It is obvious that the sizes of the cured shapes are different. Thus the input energy levels are different.

We then expose them for the same period of time. The result is shown in Figure 18. It is obvious that the sizes of the cured shapes are different. Thus the input energy levels are different. The light intensity at the center is stronger since the related cured shape is bigger.(b) pixel at the center(a) pixel around the boundaryFigure 18. Physical experiment result for energy verification.7. CONCLUSION AND FUTURE WORKMask image planning is an important process planning step for the large-area mask projection stereolithography. Optimized pixel blending is a mask image planning technique that can significantly increase the accuracy and resolution of the SFF process. To further improve its accuracy and resolution, we present a light intensity calibration system to capture the non-uniformity of a low cost off-the-shelf DLP projector. The calibration of the important pixel properties is conducted in two steps, geometricand energy calibrations.

The calibration of the important pixel properties is conducted in two steps, geometricand energy calibrations. In the geometric calibration, magnified pictures are taken and the geometric parameters of a pixel are computed based on image analysis. In the energy calibration, the light intensity was indirectly calculated by using a photosensor. Finally the integrated calibration results enable us to compute a library of light intensity functions. Such approximation functions for each pixel can correctly simulate the pixel blending results.Our future work includes the following. (1) We will incorporate the calibration results in our optimized pixel blending system. This will require us to improve our optimization solver to consider such approximation functions for each pixel. (2) We would like to investigate the large-area mask projection stereolithography based on multiple DLP projectors for better part quality.REFERENCES1. Bertsch, A., J. Jezequel, and J. Andre (1997).

libration of a Piezo-Electric Printhead in the Selective Inhibition Sintering (SIS) Process for Fabrication of High Quality Metallic PartsPayman Torabi*, Matthew Petros*, Behrokh Khoshnevis*Daniel J. Epstein Department of Industrial and Systems Engineering, Los Angeles, CA 90089Abstract Selective Inhibition Sintering (SIS) is a disruptive Additive Manufacturing process capable of printing parts from polymers, metals and ceramics. In this paper the application of a commercial piezo-electric printhead in SIS-metal is studied. This replaces the single-nozzle solenoid valve previously used in the process and allows the fabrication of high quality metallic parts due to smaller droplet sizes as well as high resolution printing mechanisms. A Design of Experiments (DoE) approach has been utilized to study the effects of important factors in printing the inhibitor. These factors include: composition of the inhibitor, quality of the print, and amount of fluid deposited for each layer.

These factors include: composition of the inhibitor, quality of the print, and amount of fluid deposited for each layer. Based on the results of these experiments, parameters have been identified for the creation of highly accurate three-dimensional parts.IntroductionAdditive manufacturing Additive manufacturing (AM) processes fabricate physical objects directly from 3D models and other digital data sources through a layer by layer manufacturing process.

There are many different processes with different methods for the fabrication of each layer of the part, but all share four common steps:i) The part is first modeled in a CAD software in computer ii) The 3D model is sliced digitally into different layers with specific thicknesses iii) The physical part is built layer by layer in the AM process iv) The part is post-processed to achieve the desired aesthetics and/or mechanicalpropertiesIndustrial processes are looking for ways to fabricate parts with higher qualities without sacrificing the production speed. With this end in mind, inkjet technologies present a tremendous opportunity to AM processes due to their high resolution and high speed print abilities. Therefore, integrating inkjet printing into a potential process can vastly increase its capabilities.Use of inkjet printheads in AM Inkjet printing involves depositing droplets of ink as small as 10 microns in diameter under precise digital control.

There are many different methods for generating and ejecting the droplets out of the print-head nozzles. Most of these methods can be classified into two main categories; continuous inkjet (CIJ) and drop-on-demand (DoD). This classification can be seen in Figure 1. Both methods print liquid through small orifices called nozzles. In CIJ the flow of the liquid is continuous. Conversely, DoD systems are impulsive, meaning small droplets are formed which can be positioned separately as needed [1].ThermalDrop‐on‐DD) Demand (DODPiezoelectricnkjet InOtherSingle jet SContinuou) us Inkjet (CIJ)s Multiple jetsFigure 1: Classsification of innkjet technologgies [1]Inkjet tec polymers form wit inkjet pri in manufchnologies c s, and chemi th appropria inting in ma facturing.can be used ical compou ate rheologic anufacturingd to print dif unds. The on cal propertie . Figure 2 sufferent types nly requirem es.

Although on the applicg [2]. inkjet there cationSelectiv SIS is a n Californi prevent s the build inhibitor next laye are repea furnace u the ones the part. steps areve Inhibitio novel AM pr ia [3]. The sintering. Fo d tank. An i is only dep er is then spr ated until al under approp treated by t An analogy summarizedg n Sintering oped by Pro rocess devel core princip ple behind S g a metallic p or fabricating ution is then nhibitor solu he boundary posited on th he previous o read over th for the part l the layers sphere. In th priate atmos . Finally, the the inhibitor would be to o think of th 3. d in Figure 3of. Behrokh K SIS is inhib part using S n deposited y of the part one and the t are comple his stage, all e inhibited r he part as if iKhoshnevis iting selecte SIS, a layer o onto select t without aff inhibitor is eted.

The pa areas of the regions are m it were insidat the Unive ed regions o of powder is ed regions o ffecting inter deposited ag art is then b e part will b manually rem de a sacrificiersity of Sou of the powd s first spread of the layer rnal regions gain. These bulk sintered be sintered e moved, reve ial mold. Thuthern der to d onto . The.

These bulk sintered be sintered e moved, reve ial mold. Thuthern der to d onto . The . The steps d in a xcept ealing e SISFigure 3: Steps forr part fabricatS tion using SISSIS-m as:metal has seeveral advanntages over oother commmercially avaailable metall processes,suchThe mach of expens The proce in other m SIS is far created w Since any itself, thehine will be sive laser or ess is faster methods, the r more scala without any re y needed su part surfacemuch cheap electron bea since only t whole area able regardin eduction in t upports are a e quality is nhe use of a c per due to th rs am generator y of the part the boundary art should be inside the pa of the printe ng the size o on of the buil the resolutio the “sacrific attached to not affected.commercial pprinthead innsteadt needs to be e scanned or ed parts.

Big ld cial mold”,e treated wh r treated [2] gger parts cahereasan beand not thee partSIS-Me Much res for the p importan researchetal Alpha search has b process. Th nt parameters is the seca Machine been conduct is is eviden s of the SIS cond generated on SIS o nced by the process [4, ation of SIover the last e quality of 5]. The SIS S-metal. Int decade resu f the printed -metal mach n the previoulting in sig d parts and hine develop ous machingnificant pro d optimizatio ped in the cu ne developeogress on of urrent d byYoozbash used for and jettin thickness generatio proof-of-hizadeh [6], depositing t ng force, th s were relati on machine) -concept fora single Do he inhibitor he penetratio ively large a and its res the processoD solenoid n onto powde on of the dr at 800 micro pective sam for metals anozzle with er layers. Du roplets into ons. In Figu mple parts ar and specificaan orifice si ue to the lack the powder ure 4, the SI re shown.

Du roplets into ons. In Figu mple parts ar and specificaan orifice si ue to the lack the powder ure 4, the SI re shown. T ally for the faize of 0.005” k of control r and conseq S-metal alph The alpha m abrication of” (0.127mm over drople quently the ha machine machine serv f bronze part) was et size layer (first ed as ts.Figuure 4: SIS-Mettal alpha macchine and some sample partts printed usinng itSIS-Me In order their corr use of a market a printhead an air bu piezoelec a drople printhead propertie the calibr materialsetal Beta M to print high responding p DOD high are commerc ds.

In a therm ubble which ctric (PZT) h t by pushin ds the bette es of the flui ration of the s for inhibitioMachine h quality par penetration resolution i cial home-us mal printhea h forces a head, a signa ng fluid out er choice fo id being prin e compositio on.rts, having c depth onto inkjet printh se photo pri d, a heated r droplet of i al is sent to t t of an orif or this resea nted such as on of the fluontrol over the powder head is need inters which resistor evap ink onto th the piezo act fice [7]. Th arch is their s viscosity, uid easier androplet sizes bed is a ne ded. The mo h utilize eith porates a sma he medium. tuators causi he greatest r higher tol surface tens nd enables ths deposited ecessity. For ost affordab her Thermal all portion o On the oth ing them to s advantage lerance for sion and den he use of dion the layer r this purpos le options i l or Piezoele of the ink for her hand, w squeeze and that makes variances in nsity.

This m ifferent candrs and se the in the ectric rming with a form PZT n the makes didateThe Epso [8], was differenton Workforc chosen and droplet sizece 30 (WF30 d incorporat es, resolution0), a high re ted into the ns and printesolution an SIS-metal t qualities thnd high perfo beta machi hat need to bormance pie ine. This pr be calibratedezoelectric pr rinthead pro d to meet therinter ovides e SISprocess r WF30 pr dedicated other threrequirements rinthead has d to black co ee color carts. The calibr s three rows olor cartridg tridges (Cyanration of the s of nozzles ges and one n, Magenta ae parameters with each row is divid and Yellow)s will be dis row contain ded into thre ) [9].scussed in la ning 180 no ee 60-nozzleater sections ozzles. Two e sections fo.

Depth n and surfache Machine f parameters print setting rs. However el of inhibit ncreases the v rint settings of penetrat ce quality ofe Paramet s that play k gs. The fluid , these two tion to occu viscosity of are importa tion determi the finishedters key roles in t d should be properties o ur, the fluid the fluid wh ant in calibr ines the lay d part.the process; e fine-tuned of the fluid d should con hich reduces rating the d yer thicknessthe compos such that i are at odds ntain a high s its jettabilit depth of pen s, which insition of the it is jettable with each o h level of so ty. Once the netration int turn affectfluid e and other. olute. fluid o the ts theInhibito The key concentra water an solution i inhibition was cond fluid that research penetrati their jettor components ation of sug d surfactant into the pow n and the jet ducted to fin t play an im provided s on into the ability withs of the inhib ar determine are importa wder.

As was ttability and nd the optim mportant role some candid powder laye the new prbitor solutio es the degre ant in the jet s previously the penetrati mum compos in its jettab date solutio er and the d rinthead. Foon are water, ee of inhibiti ttability of t mentioned, t ion of the so sition of the bility are vis on composit degree of inh or this purpo, sugar, and ion that can the solution there is a tra olution. Ther inhibitor so cosity and s tions based hibition. The ose their visan industria be achieved as well as p ade-off betw refore a serie olution. The surface tensi d on the in ese solution scosities weal surfactant d. The amou penetration o ween the degr es of experim properties o ion. Previou nhibitors eas ns were teste re measuredt. The unt of of the ree of ments of the s SIS se of ed for d andcompared viscosity determind to the visc y approxima ning the viscocosity of the ately matchin osities of theoriginal Ep ng the ink. e inhibitor sopson ink.

e inhibitor sopson ink. Th Figure 6 s olutions ande goal was t shows the e the ink.to choose th experimentahe solution w al setup usewith a d forFiguure 6: AR 20000ex viscosity tet est equipmenture of 24 oC he original E le 2.The expe strain, an ink and therimental pa nd a varying he chosen inarameters for g frequency t nhibitor solur the rheome that ranged f ution are showeter were set from 1Hz- 3 wn respectivt at a consta 35 Hz.

The r vely in Tableant temperatu results for th e 1 and TablC, 1% EpsonTable 1: VViscosity test rresults for maggenta inkFreequency 8Hz 10Hz 12HzV Viscosity 19.3 cP 16.6 cP 16.2 cPTable 2: VViscosity test rresults for inhin ibitor solutionFreequency 8Hz 10Hz 12HzV Viscosity 11.7 cP 12.3 cP 12.5 cPThe aver ink was 1rage viscosit 17.4 cP.ty of the inhiibitor solutioon is 12.2 cPP and the avverage viscosity of the EEpsonPrinter As menti and resol factors a inhibitorSettings ioned earlier lution of the along with th fluid into thr, Epson WF e prints, and he number o he powder, pF30 provide d the numbe of prints on owder surfaes different o r of nozzles n each layer ace quality afoptions for t s engaged in r determine fter droplet dthe ink drop n the printin the penetrat deposition, aplet sizes, qu ng process.

T tion depth o and print speuality These of the eed.Before calibrating the printer settings, the powder layer thickness needed to be adjusted. The metal powder used in this research was a fully alloyed bronze with chemical composition presented in Table 3. The powder particle size has a distribution of 98.7% -325mesh, 1.3% - 200/+325mesh as reported by the manufacturer. Thus, the layer thickness cannot theoretically be less than the largest particle size, approximately 80 µm. For calibration of the layer thickness, the current spreading mechanism in the machine was used to spread layers with differing thicknesses. The minimum layer thickness that could be spread consistently was 120 microns. This layer thickness was used in the determining the print settings. It should be noted that the SIS process is applicable to smaller layer thicknesses.

This layer thickness was used in the determining the print settings. It should be noted that the SIS process is applicable to smaller layer thicknesses. The limiting factor in the current study was the particle size.Table 3: Chemical Composition of the used bronze powderCopper Tin Lead Zinc Iron Phosphorous90% 10.00% 0.025% 0.04% 0.058% 0.085%A design of experiments approach was utilized here to investigate the significant and optimum values of the print parameters. The goal was to choose the best combination of factors that gives a depth of penetration of slightly more than 120 microns, an acceptable surface quality, and a satisfactory print speed. Each DoE consisted of a set of small squares printed on a thick layer of loose powder. Each square was printed with a unique combination of the controllable factors.

Each DoE consisted of a set of small squares printed on a thick layer of loose powder. Each square was printed with a unique combination of the controllable factors. The factors included: printed color which determines the number of nozzles engaged, print quality which determines the droplet sizes, and number of passes which directly affects the penetration depth as well as the surface quality. Three different levels were considered for each factor. The factors and their corresponding levels can be seen in Table 4.Table 4: Important factors and their levels used in the DOEFactor1 2 3Description Color # of Passes Print QualityLevels 1 Magenta 7 Photo0 Cyan 4 Text/Image2 Both (Blue) 10 Best PhotoThe experimental design is shown in Figure7a. Table 5 shows the print factors for the first 5 squares of the experiment. After printing the squares with their assigned factors, they were bulk sintered in the furnace.

Table 5 shows the print factors for the first 5 squares of the experiment. After printing the squares with their assigned factors, they were bulk sintered in the furnace. The inhibited powder in the printed areas was then removed (Figure7b). As can be seen in the figure, the final part consists of small squares with different depths of penetration. By measuring the depth of each square and visually comparing the surface qualities, the settings that give the most penetration depths as well as the best surface qualities were determined. The experiment was repeated a three times to ensure repeatability.Figure 7:((a) Experimenntal design (b)) The printedpartTable 5: FFactor combinnations for thefirst 5 squareesSquare NNo.ColorNo.

of Print NtsPrinnt Quality1Cyan (C)10 1Phooto (p)2Blue (B)10 1Texxt/Image (TI)3Blue (B)4 4Besst Photo (BP)4Magenta (MM)10 1Besst Photo (BP)5Cyan (C)4 4Phooto (P)Based on (which e This resu illustrate print qua slightly o speed an different . machine.n these exp equals to an ults in a fas s the penetra alities. As sh over 120 mi nd surface q print qualitperiments, it RGB value ster printing ation depth v own in the f crons as des quality are a ties. Basedt was obser of R0, G0, g process du vs. number o figure, four p sired. To ch also conside on the tablrved that pr , and B255) ue to less p of prints for prints with b oose among ered. Table le, photo qurinting with gives the g passes requir printing wit both colors p g the print q 6 summariz uality is setboth color greatest dept red for each th both color provide a dep qualities othe zes the com t as the prins simultane th of penetra h layer.

On the le gap size. Fi rinted with ththe design p les with dif eft several l nally, the sh his machineart fabricate fferent dime ines with d hapes on the under the deed to test the ensions whic different thic e right demo efined parame resolution ch show th cknesses are onstrate the m meters.of the mach he smallest h e printed th minimum whine. On the hole that ca hat determin wall thicknesse top, an be ne the s thatFiigure 9: Fabriicated part forr testing the rresolution of thhe machineThe resullts of the meeasured featuures are demmonstrated inn Table 7.Table 77: Feature size measuremenntsFeat Gap Squ Walture p size de are Hole Si s ll ThicknessMeasu 330 820 250urements (mmicrons)Conclu In this re a high re the high experime manufact beta macsion esearch an ad esolution pie potentials ents are also turing of me chine.dditive manu ezo-electric p of the SIS o carried ou etallic parts.ufacturing m printhead.

T process in ut with the m Figure 10 remachine was The develope the additive machine to t epresents sos designed an ed SIS-meta e manufactu test and ver me basic pand manufact al beta mach uring field. rify its capa arts fabricatetured that ut hine demonst The prelim ability in add ed with SIS-mtilizes trates minary ditive metalFigure 10: Baasic shapes fabricated withthe SIS-metaal processFuture The futu mechanic surface q shapes anResearch ure researche cal propertie quality of th nd generate les will be c es of the par he parts, and layer slices sconcentrated t such as the d developin specificallyd on the rob eir strength a ng slicing so for the SIS pbustness of and porosity oftware that process.the machine y, studying t can handlee, improvin the shrinkag e complicateg the e and ed 3dReferences[1] Hutchings, Ian M., and Graham D. Martin, eds. Inkjet technology for digital fabrication. John Wiley & Sons, 2012.[2] Gibson, Ian, David W.

libration of the Johnson-Cook Material Model for Additively Manufactured 304L SS Parts: Modeling and ExperimentsM. Rangapuram1, S. K. Dasari1, S.P. Isanaka1, M.F. Buchely2, J. W. Newkirk2 and K. Chandrashekhara11Department of Mechanical and Aerospace Engineering 2Department of Materials Science and Engineering Missouri University of Science and Technology, Rolla, MO 65409Abstract: Selective laser melting (SLM) is a type of additive manufacturing technique which uses a powder bed to form complex metal parts in a layer-by-layer process. This study aims to understand the material flow of parts manufactured by SLM process using 304L stainless steel powder, which is widely used in numerous applications. The tensile specimens were manufactured using 304LSS powder through SLM process. Low strain-rates, high temperature tensile tests were carried out to calibrate the parameters of the constitutive Johnson-Cook strength model.

Low strain-rates, high temperature tensile tests were carried out to calibrate the parameters of the constitutive Johnson-Cook strength model. To conduct the tensile tests, different temperatures (25 oC, 150 oC and 250 oC) and strain-rates (0.1 s-1, 0.01 s-1 and 0.001 s-1) were used. The material model developed was used in numerical simulation of the tensile tests and compared with the experimental results.1. IntroductionOwing to its design flexibility, dexterity, and reduced material loss additive manufacturing (AM) is gaining a lot of importance in recent times [1]. Laser Powder Bed Fusion (LPBF) is a type of additive manufacturing process in which a uniform powder bed resulting from a spreading process interacts with a laser source to irradiate the powder in a sequential layer process. LPBF is typically used manufacturing complex structures [2]. Historically researchers investigated the influence of build parameters on the manufactured part [3-5].

LPBF is typically used manufacturing complex structures [2]. Historically researchers investigated the influence of build parameters on the manufactured part [3-5]. Apart from the process parameters itself the quality and the properties of the printed part also depend on the type of material, the conditions used for manufacturing, machine and methodology used in the printing process. The rapid melting, solidification of the powder involves complex physical phenomena happening at the same time [6]. The rapid cooling rates which are quite inherent with the process result in non-equilibrium phases and influences the microstructure of the material. This is the reason for the difference in the properties and behavior from the conventionally manufactured counterparts [7].Some of the prominent works by researchers in characterizing the material as well as material behavior of through AM is reported below. The process parameters in the AM affects the anisotropy and mechanical properties.

The process parameters in the AM affects the anisotropy and mechanical properties. Optimizing these process parameters can improve the performance of the manufactured part. Popovich et al. [8] studied the relationship between the build orientation and the mechanical properties of the Ti6Al4V. Miranda et al. [9] assessed the properties of 316L stainless steel using a predictive model. Ortiz et al. [10] investigated the tensile properties of 304L stainless steel varying the orientation, part thickness and the build size. Guan et al. [11] studied the mechanical properties using different hatch angles.767Optimizing the parameters solely based on experimental effort can be time consuming and costly. There is a need to use the computational effort to develop models that optimize the process parameters and predict the properties [12]. Zhang et al. [13] used a thermal model to study the laser power and scan speed effects on the melt pool.

Zhang et al. [13] used a thermal model to study the laser power and scan speed effects on the melt pool. Due to the dynamic non-homogeneity of the AM processed parts the tensile and the yield strength were four times and two times higher than the as-cast material [14]. The study that involves the plastic deformation of any material will be specific to the material and the experimental effort may take a long time. This can be studied using Finite Element (FE) models. Segebade et al. [15] studied the anisotropic deformation characteristics of AlSi10Mg. The deformation mechanisms in biocompatible Ti6Al4V were investigated by Kadkhodapour et al. [16]. Their study revealed that the use of Johnson-Cook damage model can estimate the stresses within 18% error. There is very limited to no data on the JC parameters for additively manufactured steels.

Their study revealed that the use of Johnson-Cook damage model can estimate the stresses within 18% error. There is very limited to no data on the JC parameters for additively manufactured steels. The material used in this study is 304L stainless steel.304L stainless steel has gained importance over the years due to its mechanical properties and chemical compositions. LPBF of 304L stainless steel to fabricate parts does not require any post processing like solution annealing due to low carbon content which reduces the precipitation of carbide [17]. 304L stainless steel manufactured through AM results in superior mechanical (tensile and yield strength) than conventionally manufactured counterparts. This makes them a perfect candidate for high strength applications [18].Johnson-Cook (JC) model is a widely used analytical model to predict the material flow at different temperatures and strain-rates [19].

These obtained parameters were used in developing an FE model to simulate the tensile behavior of the AM material. The accuracy of the developed model was compared to the actual tensile testing experiments.2. Materials and methods2.1 Manufacturing of 304L SS with LBPF768The tensile specimens were manufactured using gas atomized 304L stainless steel powder with a particle size ranging from 15 µm to 63 µm. The chemical composition (wt%) of the material can be found in Table 1. The fabrication of the 304L SS parts was carried out using Renishaw AM250 machine. The machine process parameters used in this study are listed below (Table 2).Table 1. Chemical composition of 304L stainless steel powderFeCNSiNiCr Mn Cu MoP69.9 0.025 0.07 0.015 7.9 17.7 1.75 0.84 0.32 0.03Table 2.

Chemical composition of 304L stainless steel powderFeCNSiNiCr Mn Cu MoP69.9 0.025 0.07 0.015 7.9 17.7 1.75 0.84 0.32 0.03Table 2. Machine parameters of Renishaw AM250Process parameter Effective build volume (mm x mm x mm) Laser power (W) Laser type Laser spot Hatch distance (mm) Layer thickness (µm) Point distance (µm) Exposure time (µs) Point distance (µm) Inert gas during productionValue 248 x 248 x 280 200 1070 nm NdYAG 70 0.085 50 60 75 60 Argon2.2 Manufacturing of tensile couponsThe Renishaw AM250 machine was used to manufacture the tensile coupons for this study. ASTM-E8 subsize specimens were printed in both XY and Z directions. The ASTM-E8 specimen with the dimensions can be found in Figure 1. The specimen was 2 mm thick with a 2o taper included in the design to induce the failure in gauge section of the specimen.Figure 1. ASTM-E8 subsize specimen (dimensions in mm)3.

The specimen was 2 mm thick with a 2o taper included in the design to induce the failure in gauge section of the specimen.Figure 1. ASTM-E8 subsize specimen (dimensions in mm)3. Low strain-rates tensile testing769The JC model was used to investigate the plastic deformation of 304L stainless steel in this work. To obtain the JC parameters of the material, low strain-rate tensile tests were performed at various temperatures and strain rates. The temperatures of 25ºC, 150ºC, 250ºC and the strain rates of 0.1 s-1, 0.01 s-1, 0.001 s-1 were chosen for this work. These ranges were mainly selected with respect to the structural application of the material. The complete experimental test matrix can be found in Table 3. Five replications were performed for each combination of temperature and strain rate.Table 3.

The complete experimental test matrix can be found in Table 3. Five replications were performed for each combination of temperature and strain rate.Table 3. Experimental test matrixStrain rate0.1 s-1 0.01 s-1 0.001 s-125 oC (0.1,25) (0.01, 25) (0.001, 25)Temperature 150 oC (0.1,150) (0.01, 150) (0.001, 25)250 oC (0.1, 250) (0.01, 250) (0.001, 250)3.1 TestingThe tensile tests were performed on an MTS frame with environmental chamber to control the temperature. The MTS machine (Figure 2a) and the specimen installed in the environmental chamber can be found in Figure 2b.(a) (b)Figure 2. (a) MTS frame with environmental chamber (b) Specimen installed in the environmental chamber4. Finite Element Model770Numerical modeling method was used to verify the data obtained from the experimental tensile tests. For this purpose, similar experimental conditions were modeled in ABAQUS using dynamic explicit analysis.

For this purpose, similar experimental conditions were modeled in ABAQUS using dynamic explicit analysis. The JC parameters obtained from the tests were used for the material model in these simulations. A number of 5040 hexagonal elements, C3D8RT type were used in meshing the model. The tensile coupon meshed in ABAQUS with the boundary conditions can be found in Figure 3. The experimental and the FEA model outcomes were compared to verify the JC parameters for the 304L SS steel alloy.Figure 3. The ASTM-E8 tensile coupon and the meshed tensile coupon in ABAQUS5. Results and discussion5.1 Stress-strain curvesFrom the tensile tests performed at different temperatures and strain rates, the following true stress- strain curves were obtained, in both XY and Z print directions (Figures 4 and 5).771Figure 4. True stress-strain curves in XY-direction (a) at 25ºC (b) 150ºC (c) 250ºCFigure 5.

True stress-strain curves in XY-direction (a) at 25ºC (b) 150ºC (c) 250ºCFigure 5. True stress-strain curves in Z-direction (a) at 25ºC (b) 150ºC (c) 250ºCFrom the XY stress-strain curves it was evident that the yield stress of the material was highest at 25 oC with 0.1 s-1 strain rate, this was observed to decrease with strain rates to 0.01 s-1 and 0.001 s- 1 at 25 oC. Similar trend was observed at 150 oC tests with highest yield stress recorded at 0.1 s-1 strain rate followed by 0.01 s-1 and 0.001 s-1. The yield stress values dropped significantly for the 250 oC tests with 0.1 s-1 strain rate recording the highest value. For the Z direction testing, the yield stress of the material was highest at 25 oC with 0.1 s-1 strain rate followed by 0.01 s-1 and 0.001 s- 1. As the temperature increased to 150 oC and 250 oC, these values were further reduced.

As the temperature increased to 150 oC and 250 oC, these values were further reduced. In772comparison between the XY and Z direction tests, the XY specimens recorded highest yield stress values at the respective temperatures and strain rates.5.2 Johnson-Cook parameters from the tensile testsA Genetic Algorithm (GA) was written to optimize the experimental data and fit it to the JC material model equation. This code utilizes all the experimental tensile testing data at different temperatures and strain rates to give us the JC parameters as well as fits these values to the equation. Details of the model can be found in Buchely et al. [20]. Thus, the JC model parameters obtained from this study can be found in Tables 3 and 4 for XY and Z direction respectively. The R2 value of the fitting was higher than 96% which signifies it to be a very good fit to the data. The experimental and the predicted values fit can be found in Figures 6.Table 4.

The R2 value of the fitting was higher than 96% which signifies it to be a very good fit to the data. The experimental and the predicted values fit can be found in Figures 6.Table 4. JC Model parameters for 304L SS in XY build directionMaterial 304L SSA (MPa) 516.45B (MPa) 812.39n 0.7042C 0.0117m 0.7383(s-1)0.001 𝜺𝜺̇ 𝟎𝟎Table 5. JC Model parameters for 304L SS in Z build directionMaterial 304L SSA (MPa) 491.83B (MPa) 571.40n 0.6422C 0.0118m 0.7727(s-1)0.001𝜺𝜺̇ 𝟎𝟎Figure 6. The experimental vs predicted values from the GA model for (a) XY direction (b) Z direction5.3 Comparison of experimental and modeling resultsFigure 7 shows the dimensions of necking region at failure from experiments at 25 oC for XY direction specimen and it is compared to FE model. It is observed that the strain at failure in FE model is same as that of experiment (Figure 4).

It is observed that the strain at failure in FE model is same as that of experiment (Figure 4). Figure 8 shows the comparison of the experimental773and modeling results for the XY and Z direction specimens at 25 oC. The predicted yield stress for both XY and Z FE models was less than 10 percent when compared to experiment.Figure 7. The experimental vs modeling comparison of XY direction specimen at 25 oCFigure 8. The comparison of the experimental and modeling results of XY and Z direction samples at 25 oC7746. ConclusionsIn summary, 304L stainless steel was used to additively manufacture the tensile coupons. From tensile tests at different temperatures and strain rates, a constitutive JC material model was calibrated using genetic algorithm. The obtained JC parameters were used in developing an FEA model to simulate the experimental tensile tests.

The obtained JC parameters were used in developing an FEA model to simulate the experimental tensile tests. This model was used to compare the experimental results and modeling results of the tensile tests to verify the developed material model. The difference in the FEA model results and the experimental results was within 10% error.References:1. S. Gorsse, C. Hutchinson, M. Gouné, and R. Banerjee, “Additive manufacturing of metals: A brief review of the characteristic microstructures and properties of steels, Ti-6Al-4V and high-entropy alloys,” Science and Technology of Advanced Materials, Volume 18, Pages 584-610, 2017.2. Y. Zhai, D.A. Lados, and J.L. LaGoy, “Additive manufacturing: making imagination the major limitation,” The Journal of The Minerals, Metals & Materials Society, Volume 66 Pages 808-816, 2014.3. E. Liverani, S. Toschi, L. Ceschini, and A.

M-LEM PROCESSING: MATERIALS FLEXIBILITYz. E. Liu, T. C. Ko, J. Best, 1. D. Cawley, and A. H. Heuer Department of Materials Science and Engineering Case Western Reserve University 10900 Euclid Avenue Cleveland OH 44106-7204ABSTRACTThe cut-then-stack paradigm of computer-aided manufacturing of laminated engineering materials (CAM-LEM) offers choice of feedstock material, ease of handling finely divided (and therefore sinterable) powders, and the ability to mix materials. This combination of features was exploited to process fluidic devices. CAM-LEM processing was used to render the Part in aluminum oxide, silicon nitride, and stainless steel.INTRODUCTIONComputer-aided manufacturing of laminated engineering materials (CAM-LEM) has beendeveloped as a solid freeform process to allow direct layered-manufacturing of components of complex geometry in anyone of an array of engineering materials.

To date, efforts has focussed on the use of powder-based processing of ceramics and metals.When working in ceramics or metals using the CAM-LEM process, the method of partconstruction is robotic assembly of cut outlines followed by green-state lamination and conventional firing (1-4]. The feedstock for the process is a "green taPe" that serves as a powder carrier. Separate preprocessing of the taPe by methods such as tape casting [5] or compression molding [6] allows the process to employ finely divided powders (submicron, in the case of ceramics), creates uniform particle packing, and completely avoids problems associated with flow instabilities or segregation effects.Powder-based taPes are widely used in industrial applications and a number of techniqueshave been develOPed for their production. SPecific process selection is dePendent on proPerty requirements. For CAM-LEM processing, important taPe characteristics include behavior during laser cutting and ease of lamination.

SPecific process selection is dePendent on proPerty requirements. For CAM-LEM processing, important taPe characteristics include behavior during laser cutting and ease of lamination. It has been determined that some commercially available taPe formulations are suitable for CAM-LEM processing, whereas others are not. In addition, taPes can be produced from either commercially compounded feedstocks or mixtures of raw materials.Through the use of the cut-then-stack paradigm, CAM-LEM processing offers increasedgeometric flexibility relative to other forms of laminated object manufacturing. The ability to carry out "tangent cutting" yields improved surface finish and faithfulness in approximation of a CAD model [7-9]. Furthermore, the ready inclusion of a fugitive material during building allows parts to by built with highly complex interior surfaces [3,4].DEFINITION OF TEST PARTOne class of parts that are well suited to CAM-LEM processing are fluidics.

These devicesare used for a variety of pUfPQses, but share a common OPerational characteristic; all oPerate by motion of a condensed fluid through a series of internal passages within the device under a pressure gradient. CAM-LEM processing is an appealing method for fabricating these parts because it allows integral fabrication (i.e., formation of a single monolithic piece with appropriate internal passages), thus mitigating the need separate processing of subcomponents and assembly, and increasing reliability. CAM-LEM also offers increased flexibility in material choice. The379material choice. The particular fluidic circuit selected for this study requires five layers, has a circular motif, and is disc-shaped with three fold rotational symmetry.MATERIALS: SELECTION, PREPROCESSING, AND LAMINATIONThree materials were selected for construction: aluminum oxide, silicon nitride, and stainless steel.

Silicon nitride powder (GS-44, AlliedSignal Ceramic Components, Torrance CA) was tape cast using a binder system based on polyvinyl butyral. Commercial (Coors Electronic Ceramics, Chattanooga) aluminum oxide tapes were used. These tapes were 96% alumina and fabricated by tape casting using a binder system also based on polyvinyl butyral. Stainless steel tapes were fabricated by compression molding using a developmental injection molding feedstock (Rohm and Haas, Springhouse PA). The steel powder was gas-atomized 316L and binder system in this case was determined to be based on polymethyl methacrylate. For working with both the silicon nitride and aluminum oxide tapes, a fugitive tape system based on graphite powder has been developed in-house. Characteristics for all tapes are given in Table 1.Table 1. Green Tape CharacteristicsTape Aluminum Oxide Silicon Nitride Stainless Steel Graphite*Thickness 600 pm 300 600 300 or 600Solids Loading 56.6 vol.

Green Tape CharacteristicsTape Aluminum Oxide Silicon Nitride Stainless Steel Graphite*Thickness 600 pm 300 600 300 or 600Solids Loading 56.6 vol. % 51.6 64.0 42.8Binder Content 20.5 vol. % 29.9 36.0 21.9Porosity 22.9% 18.5 none 35.3Thicknesses were processed to match various feedstocks. For this system. all calculations assume a graphite density of 2.3 g/cc.For all systems, lamination was achieved using a variant of the adhesive lamination processthat has been reported elsewhere [10]. The specifics of the adhesive formulation were varied for each system in order to optimize lamination efficiency and to obtain process tolerance (e. g., allow good adhesion throughout a broader window in time after the adhesive formulation was applied). Both the nature of the organic "package" and the characteristics of the porosity influence the choice of adhesive.CAM-LEM PROCESSED PARTSSuccessful parts were produced from all three materials.

Both the nature of the organic "package" and the characteristics of the porosity influence the choice of adhesive.CAM-LEM PROCESSED PARTSSuccessful parts were produced from all three materials. Figure 1 shows the cut outlines for both the aluminum oxide and graphite fugitive. Figure 2a shows an assembled part made from aluminum oxide tape with a graphite fugitive. The fugitive plays two roles; it defines the internal passages and it serves as a pressure-transmitting medium during lamination. Figure 2b is an image of the same part after pyrolysis of the binder and removal of the fugitive have been effected. Note that the fugitive is always removed prior to the onset of macroscopic shrinkage. This is done to avoid the occurrence of backstresses or encapsulation of fugitive by a densifying matrix.

Note that the fugitive is always removed prior to the onset of macroscopic shrinkage. This is done to avoid the occurrence of backstresses or encapsulation of fugitive by a densifying matrix. Figure 2c is an image of the fully densified part, in which the nominally 17% linear shrinkage is observed.Figure 3a and 3b are backlit images of the translucent aluminum oxide part from two sides, respectively. The interior passages are clearly imaged and these figures demonstrate that the details of the interior surface are well preserved during postprocessing, i.e., firing.Similarly good results were obtained using silicon nitride. Figure 4 shows an image of oneface of a completely processed part. The tearshaped contrast is produced by low-level carbon doping of the silicon nitride in the outer layer where it was in contact with the fugitive underneath. This has been shown not to affect the properties of the material [11].

This has been shown not to affect the properties of the material [11]. As silicon nitride is opaque,380it was not possible to image the interior directly, but functionality was demonstrated by simple flow tests.The stainless steel version of this part is shown in Figure Sa and 5b. A high quality partalso was produced in this material.High lamination efficiency was confirmed by sectioning and preparing the cut surfaces using standard metallographic techniques. External dimensions of aluminum oxide and silicon nitride were determined using a coordinate measuring machine. Line scans across the diameter of the face containing the larger oriface were collected from several aluminum oxide samples as well as one stainless steel, and a pseudo-2D map was generated of the same surface of a silicon nitride part.

This surface was chosen, rather than the opposite, because the three teardrop-shaped voids in the second layer present a more critical test of the ability to process parts with interior channels.The line scans from the alumina indicate that the inclusion of the fugitive does not significantly affect the tendency of the sample to slump during firing. Line scans revealed that a small amount of slumping, on the order of 150}tm over a 15 mm span, occurred over internal cavities in all cases. A typical result is shown in Fig. 6. Similarly, the surface relief map of the silicon nitride part shows modest relief, see Fig. 7. The line scans from stainless steel part also show evidence of relief, but it is considerably more modest.

7. The line scans from stainless steel part also show evidence of relief, but it is considerably more modest. The difference is attributed to the fact that both of the ceramics form an appreciable amount of liquid during firing (due to the presence of silicate sintering aids) whereas under the conditions used to sinter the stainless steel, the system remains completely solid. The results are entirely consistent with the expectations associated with these materials; the deformation during firing is comParable to that observed in these materials when conventionally processed.CONCLUSIONSThe use of CAM-LEM processing to directly produce a component in a variety of structuralmaterials has been demonstrated. Thus, CAM-LEM processing is an attractive method for conducting small scale production in the context of materials substitution trials.A fugitive has been developed that is compatible with both aluminum oxide and siliconnitride.

The fugitive has been used to produce integral parts with highly complex internal geometrical features.ACKNOWLEDGEMENTSThis research was supfX>rted by the CAMP, Inc. Applied Research Program and the Office of Naval Research (ONR Grant Number NOOOI4-95-10107). Aluminum oxide tape was supplied by Dr. Debra Hom of Coors Electronic Ceramics. Dr. John Pollinger, AlliedSignal Ceramic Components, supplied the silicon nitride powder and fired the silicon nitride parts. The stainless steel feedstock was supplied by Dr. Len Bogan of Rohm and Haas. Firing of the stainless steel parts was carried out at BFGoodrich, Brecksville OH, by Tom Nixon. Dr. Ajit Sane granted access to compression molding facilities at BP America Finally, CAM-LEM is a multidisciplinary activity at Case Western Reserve University, we gratefully recognize the contributions of our colleagues in the Electrical Engineering and Applied Physics Department; Prof. W. S. Newman, Dr. Y. Zheng, and S. Choi.REFERENCES1. J. D.

6th Eur. Conf. on Rapid Prototyping and Manuf., Univ. Nottingham, UK, pgs. 21-30, (1997).10. P. Wei, 1. D. Cawley, and A. H. Heuer, "CAM-LEM Processing: Lamination Technology," Proc. 6th Eur. Conf. on Rapid Prototyping and Manuf., Univ. Nottngham, UK, pgs. 95-104 (1997).11. Personal communication, J. Pollinger, AlliedSignal Ceramic Components.Figure 1 Digitial photographs of the cut outlines forming the positive space (from alumina, upPer series) and the negative space (from fugitive, lower series) of the fluidic circuit selected for study.382Figure 2 Aluminum oxide fluidic circuit after assembly.

pillary-Driven Flow in Open Microchannels Printed with Fused Deposition ModelingRobert K. Lade Jr.1, Erik J. Hippchen1, Luke Rodgers2, Christopher W. Macosko1, and Lorraine F. Francis11Department of Chemical Engineering and Materials Science, University of Minnesota 421 Washington Ave. SE, Minneapolis MN, 55455, USA2Stratasys Ltd., 7665 Commerce Way, Eden Prairie, MN, 55344, USAAbstractThe fundamentals of fluid flow in 3D printed, open microchannels created using fused deposition modeling (FDM) are explored. Printed microchannels are used in microfluidic devices and have potential applications in embedding electronics in plastic substrates. However, FDM parts possess rough surfaces, and in this study, surface topography is shown to have an important impact on flow behavior, causing the liquid to travel down the channel with a characteristic ‘pulsing’ movement.

We also analyze the influence of print orientation on capillary flow, where microchannels printed in specific orientations are shown to exhibit different flow dynamics.Keywords: Fused deposition modeling ; capillary flow ; microchannel ; flow dynamics1.IntroductionFluids are brought into contact with three-dimensional (3D) printed parts for a variety of reasons. For example, liquid coatings are often applied after printing to smooth,1 seal,2 strengthen, or add functionality3 to a part. Additionally, 3D printed microfluidic devices have been used by several researchers4–6 for a variety of applications; all of these devices involve interfacing fluids with a 3D printed surface. While 3D printing microfluidic devices has become common,7 little is known about how the 3D printed surface influences the flow dynamics.

While 3D printing microfluidic devices has become common,7 little is known about how the 3D printed surface influences the flow dynamics. Because many 3D printing technologies produce parts with rough surfaces,8–10 the flow dynamics can potentially be more complicated than on flat surfaces.The increasing number of studies involving direct contact between fluids and 3D printed surfaces merits a more in-depth understanding of the physical principles and dynamics governing this interaction. In this paper, we focus on the fundamentals of fluid flow on parts printed with fused deposition modeling (FDM). Specifically, we look at the behavior of capillary flow in FDM- printed open microchannels.

Specifically, we look at the behavior of capillary flow in FDM- printed open microchannels. Capillary flow in 3D printed microchannels not only has important applications in microfluidic device design, but also has potential applications in embedding electronics in 3D printed parts.11 FDM was chosen as an initial 3D printing technology for study due to its flexibility and accessibility.2.Experimental2a. Microchannel design and printingAll microchannels were printed using a Dimension 1200es 3D FDM printer (Stratasys, Inc., Eden Prairie, MN). Acrylonitrile butadiene styrene (ABS P-400, Stratasys, Inc.) was used as the model material and a soluble support (P-400SR, Stratasys, Inc.) was used as the supportmaterial. All microchannels were printed with a layer resolution of 254 µm (0.01 in.) with solid interior fill and default temperature settings.A cartoon of one of the CAD designs is shown in Fig. 1a and a printed capillary filled with liquid is shown in Fig. 1b.

1a and a printed capillary filled with liquid is shown in Fig. 1b. Each capillary consists of one reservoir connected to a channel, with widths ranging from 80 - 750 µm and depths of 250 µm. Reservoir diameters were designed such that the reservoir and channel volumes were equal.(a)(b)Figure 1. (a) Cartoon of microchannel CAD design. Not to scale. (b) Coated capillary printed from a CAD design specifying a 400 µm wide channel. Channel is filled with water with red dye. Scale bar: 1 cm.Each part was printed in one of three orientations (Fig. 2a). Because FDM employs cylindrical strands of material, FDM parts have a characteristic rough surface. Accordingly, each orientation imparts a unique topography to the bottom of the microchannel. Capillaries oriented at 45° and 135° have channels which are oriented parallel with and perpendicular to the printed layer beneath it, respectively. Figs.

Capillaries oriented at 45° and 135° have channels which are oriented parallel with and perpendicular to the printed layer beneath it, respectively. Figs. 2b-d shows cross sections (parallel to the length of the channel) of microchannels printed in each orientation. The microchannel geometries illustrated in Figs. 2b, c, and d will henceforth be referred to as diagonal, parallel, and perpendicular orientations, respectively.(b)(a)(c)(d)Figure 2. (a) Cartoon of microchannel part rotated 0°, 45°, and 135° on build stage. During printing, the print head rasters back and forth at 45° in the x-y plane. Cross sections (parallel to the length of the channel) of microchannels printed in (b) diagonal (0°), (c) parallel (45°), and (d) perpendicular (135°) orientations. Scale bars: 200 µm.2b. Microchannel surface treatmentAfter printing, microchannels were coated with an ultraviolet (UV) curable adhesive (NOA73, Norland Products, Inc., Cranbury, NJ) to make the channel watertight.

Microchannel surface treatmentAfter printing, microchannels were coated with an ultraviolet (UV) curable adhesive (NOA73, Norland Products, Inc., Cranbury, NJ) to make the channel watertight. Before coating, each capillary was washed with water, air dried, and plasma treated. After plasma treatment, a foam brush (model no. 8500 0100, Home Depot, Atlanta, GA) was used to apply the coating over the entire part. The adhesive was then cured under long-wave UV light (366 nm) for 45-60 minutes.2c. Monitoring capillary flowCapillary flow was monitored using a digital optical microscope (KH-7700, Hirox USA, Hackensack, NJ) at 50x zoom. The test liquid was an aqueous solution of 84 wt. % glycerol mixed with 40 µL of blue food dye per 20 mL of glycerol solution. The viscosity and surface tension of the test solution was ~80 mPa∙s and 62 mN/m, respectively, and its contact angle with the coated ABS was ~40°.

The viscosity and surface tension of the test solution was ~80 mPa∙s and 62 mN/m, respectively, and its contact angle with the coated ABS was ~40°. Prior to each experiment, each capillary was plasma treated to encourage wetting and spreading. After dispensing the glycerol solution into the reservoir, videos of the flow were recorded and the position of the capillary front in each frame was extracted using ImageJ software (v1.48v, NIH).3.Results and Discussion3a. Microchannel geometry and surface topographyA sample cross section of a microchannel is shown in Fig. 3, taken perpendicular to the length of the channel. Microchannel walls are formed by individual printed strands, giving each microchannel a characteristic rounded shape. Although print orientation does not affect the shape of the microchannel, it does affect print accuracy.

Although print orientation does not affect the shape of the microchannel, it does affect print accuracy. Microchannel dimensions were determined using optical microscopy and, in general, microchannels printed in a parallel orientation match their CAD designs most closely. In all orientations, wider microchannels are printed more accurately. Microchannel depth was relatively consistent across all orientations, with an average value of 260 ± 30 µm.Figure 3. Cross section of FDM microchannel, perpendicular to the length of the channel. Scale bar: 200 µm.The surface topography of the microchannels printed in each orientation is shown in Figs. 2b-d. While both the diagonal and perpendicular print orientations present the liquid with relatively large topographical obstacles during flow down the channel, the parallel orientation presents a relatively smooth surface. The impact that these topographies have on the capillary flow will be discussed in the next two sections.3b.

The impact that these topographies have on the capillary flow will be discussed in the next two sections.3b. Capillary flow dynamicsThe FDM microchannels explored in this study exhibited unique flow dynamics. Fig. 4 shows a representative plot of meniscus (i.e., capillary front) position vs time for a microchannel of width ~300 µm printed in a diagonal orientation. The most distinguishing feature of this flow is the small jump in position every 0.5 - 1s. Instead of flowing smoothly down the channel, the flow exhibits a characteristic “pulsing” motion as a result of the surface topography. As the meniscus approaches a trough between two printed strands (see Fig. 2b), it becomes pinned for a short time. Once that pinning is overcome, the meniscus then surges forward quickly due to this build up liquid, with the edge that became pinned first becoming unpinned first.Figure 4. Meniscus position vs time for a microchannel with width of 320 µm, printed in a diagonal orientation.

Meniscus position vs time for a microchannel with width of 320 µm, printed in a diagonal orientation. Jumps in position occur every 0.5 – 1 s. Arrows indicate some of these jumps.This pulsing flow continues down the entire channel, with pinning time varying from trough to trough, corresponding to small variations that arise during printing. For example, in Fig. 4 it is clear that after about 0.5 s, the liquid became pinned for a relatively long time (~ 1 s) before it started to flow again whereas the pinning visible at 2 s lasts less than 0.5 s. Arrows indicate the jumps that occur before and after the pinning around 1 s. Thinner capillaries were observed to exhibit shorter pinning times and hence an overall smoother flow, while in larger capillaries the pulsing flow was more pronounced.Effect of print orientation Capillaries printed in a perpendicular orientation exhibit a strong, uniform pulsing flow.

This is due to the topography being completely perpendicular to the flow such that neither the top nor the bottom of the meniscus becomes unpinned first. This feature results in a slightly longer pinning of the meniscus and a proportionally faster pulse upon depinning. Oppositely, microchannels printed in a parallel orientation exhibit no pulsing. Because there are essentially no topographical barriers along the length of the channel (see Fig. 2c), the flow can proceed uninterrupted.3c.Despite the vastly different flow dynamics that each capillary orientation imparts to the liquid flowing down it, the average measured velocity was observed to be essentially independent of print orientation. This result suggests that the surges that occur at each trough on the diagonallyand perpendicularly oriented capillaries offset the slowing down caused by the pinning essentially equally, resulting in no net effect on the average velocity as it moves down the microchannel.

Flow velocity was observed to increase as channel width decreased. At the smallest printable channel size (~200 µm), flow velocities around 1000 µm/s were recorded. At the largest widths tested (~700 µm), flow velocities around 200 µm/s were recorded.4.Summary and future workIn this work, we showed that due to the nature of the FDM printing process, printed microchannels have a topography that depends on orientation. This topography causes the capillary flow to move with a pulsating motion, resulting from a constant pinning and depinning at each topographical feature. Despite this unique motion, the average velocity of the flow is not affected by differences in the topography. When printing microchannels, the minimum feature sizes are set by the individual capabilities of the printer.

When printing microchannels, the minimum feature sizes are set by the individual capabilities of the printer. Here, we were able to achieve minimum channel depths and widths of approximately 250 µm and 200 µm, respectively.Future work will focus more on a quantitative understanding of the flow dynamics in each differently oriented microchannel. Although FDM is a relatively simple and accessible printing technology, other technologies with higher resolution may offer better opportunities for printing microchannels. Future work will focus on comparing the FDM microchannels in this study to microchannels printed using stereolithography.

Future work will focus on comparing the FDM microchannels in this study to microchannels printed using stereolithography. This work provides a foundation for developing a more in-depth understanding of the fundamentals governing flow in 3D printed microfluidic devices and also promises to promote the development of other applications involving microchannels embedded in 3D printed parts.5.AcknowledgementsThe authors would like to thank the industrial supporters of the Coating Process Fundamentals Program (CPFP) of the Industrial Partnership for Research in Interfacial and Materials Engineering (IPRIME) for supporting this research. This work was also supported by funding through MnDRIVE. Parts of this work were carried out at the Minnesota Nano Center, which receives partial support from NSF through the NNIN program.

This work was also supported by funding through MnDRIVE. Parts of this work were carried out at the Minnesota Nano Center, which receives partial support from NSF through the NNIN program. The authors would like to extend their gratitude to Satish Kumar at the University of Minnesota for helpful discussions.6.References(1) Zhu, J.; Chen, J.; Lade Jr., R.K.; Suszynski, W.J.; Francis, L.F. Water-Based Coatings for 3D Printed Parts. J. Coatings Technol. Res. 2015. doi: 10.1007/s11998-015-9710-3.(2) Mireles, J.; Adame, A.; Espalin, D.; Medina, F.; Winker, R.; Hoppe, T.; Zinniel, B.; Wicker, R. Analysis of Sealing Methods for FDM-Fabricated Parts. Technical report, W.M. Keck Center for 3D Innovation. The Univ. of Texas, El Paso (2011).(3) Zhu, J. Water-Based Antimicrobial Coating for 3D Printed Medical Devices. Stratasys White Paper.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCarbide Formation in Nickel-Base Superalloy MAR-M247 Processed Through Scanning Laser Epitaxy (SLE)Amrita Basak1, and Suman Das1, 21George W. Woodruff School of Mechanical Engineering, 813 Ferst Dr NW, Georgia Institute of Technology, Atlanta, GA 30332 2School of Materials Science and Engineering, Georgia Institute of Technology, 771 Ferst Dr NW, Atlanta, GA 30313AbstractNickel-base superalloys develop high-temperature strength primarily due to the solid-solution- strengthening and the precipitation-strengthening mechanisms typically through cobalt/chromium and aluminum/titanium respectively. Certain other elements such as boron and zirconium are chosen for grain boundary strengthening.

Certain other elements such as boron and zirconium are chosen for grain boundary strengthening. Such elements tend to segregate to the grain boundaries reducing the grain boundary energy and resulting in better grain boundary cohesion and ductility. Another form of grain boundary strengthening is achieved through the addition of carbon and various carbide formers. The carbide formers are responsible for driving precipitation of carbides at grain boundaries and thereby reducing grain boundary sliding. Various types of carbides such as blocky, elongated, and Chinese-script are possible in the microstructures of nickel-base superalloys depending on the composition of the superalloy and processing conditions. However, in the SLE fabricated MAR-M247, only blocky carbides were predominantly observed. Scanning electron microscopy and energy dispersive X-ray spectroscopy investigations were carried out and the carbides were found to be tantalum-rich.

Scanning electron microscopy and energy dispersive X-ray spectroscopy investigations were carried out and the carbides were found to be tantalum-rich. This work is sponsored by the ONR through grant N00014-14-1-0658.IntroductionNickel-base superalloys are extensively used in aircraft and power-generation turbines, rocket engines, and other challenging high-temperature environments as these alloys offer higher yield strength with increasing operating temperatures primarily due to the presence of the secondary  phases that form an antiphase boundary and lock the associated dislocation movements. For the last few decades, extensive research in superalloy development has culminated in productionization of components with alloys that can withstand temperatures up to 90% of the melting point of the material [1].

However, precipitation of the  phases causes strain age cracking; hardening of the alloy during thermal exposure, and transfer of the solidification strains onto the grain boundaries. The susceptibility to strain age cracking and the alloy weldability is often qualitatively described for nickel-base superalloys by plotting aluminum (Al) vs. titanium (Ti) content in the alloys. The alloy is considered difficult-to-weld if the total Al + Ti content exceeds 4 wt. % [2]. The total Al + Ti content for polycrystalline MAR-M247 is 6.0 wt. %, thus making it a difficult-to-weld alloy.The degradation of the hot-section components is attributed to creep, thermo-mechanical fatigue, high-temperature oxidation, and thermal corrosion. Currently, the lack of suitable commercial repair technologies results in high maintenance costs for these components that can460reach as much as hundreds of thousands to even millions of dollars per engine.

Hence, there is a great commercial interest in developing additive manufacturing (AM) processes that can restore the parent metallurgical microstructure and geometry at damage locations for these parts. Scanning laser epitaxy (SLE) is a metal powder bed fusion-based additive manufacturing process that creates equiaxed (EQ), directionally solidified (DS) and single-crystal (SX) structures in nickel-base superalloys through controlled laser melting of alloy powders onto like chemistry substrates. In SLE, a galvanometer controlled high power laser beam is directed onto the powder bed. Processing parameters such as laser power, scan spacing, and scan speed are varied in SLE to control and optimize the deposit characteristics. An important criterion for a successful SLE deposit is the determination of energy density that ensures a good metallurgical bond between the substrate and the deposit. The metallurgical bond starts at one edge of the substrate where the melt pool is initiated.

The metallurgical bond starts at one edge of the substrate where the melt pool is initiated. The melt pool travels along the length of the substrate fusing the powder to the substrate. SLE has been successfully demonstrated to deposit SX CMSX-4 [3-5], EQ René 80 [6] and EQ IN100 [7] on similar chemistry substrates.MAR-M247 is an equiaxed (EQ) or polycrystalline nickel-base superalloy that was developed at the Martin Marietta Company in the 1970s [8]. Over the last few decades, this alloy has been extensively used in the production of advanced turbine blades and rotating parts in the aerospace industries because of its excellent thermo-mechanical properties. The alloy demonstrates high creep strength and good castability along with excellent oxidation and corrosion resistance at elevated temperatures. MAR-M247 is a nickel-base, vacuum melted, cast superalloy with a high  [Ni3 (Al, Ti)] volume fraction (62%). The alloy also contains a significant amount (13.7 wt.

MAR-M247 is commonly used for blade rings and high-pressure turbine blades in gas turbine engines.The major constituents of the SLE deposited microstructure were the  matrix, the  precipitates in the  matrix, the eutectics, and the carbides. In the present study, morphologies of the carbide precipitates in the SLE deposited MAR-M247 were characterized. In the SLE deposited MAR-M247, the carbides precipitated in the interdendritic region. The investment cast substrate region revealed the presence of blocky, elongated and Chinese-script carbides. However, in the deposit region, blocky carbides were predominantly present as revealed by the scanning electron microscopy (SEM) investigations. The interface region shows the presence of a few elongated carbides on the deposit side. Energy dispersive X-ray spectroscopy (EDS) measurements revealed that the blocky carbides are mostly Ta-rich. Carbide precipitates near the interface showed trace amounts of boron.

Energy dispersive X-ray spectroscopy (EDS) measurements revealed that the blocky carbides are mostly Ta-rich. Carbide precipitates near the interface showed trace amounts of boron. Commercially available heat treatment was carried out and the results showed the formation of Cr-rich plate-like discrete carbide network at the grain boundaries. However, the morphologies of the blocky carbides remained unchanged during the heat treatment.461Materials and MethodsPowder of nickel-base superalloy MAR-M247 produced by the Praxair Surface Technologies through Argon gas atomization was used in this study. The composition of the powder is reported in Table I. The morphology and the cross-sections of the powder particles were analyzed using optical microscopy and scanning electron microscopy (SEM). The MAR-M247 powders were mostly spherical, although some irregularity in shape was found as shown in Fig. 1(a). The powder was mounted in Bakelite and polished to a mirror finish.

The MAR-M247 powders were mostly spherical, although some irregularity in shape was found as shown in Fig. 1(a). The powder was mounted in Bakelite and polished to a mirror finish. The cross-section was analyzed under an optical microscope for the inspection of internal porosity, and no major internal porosity was detected as shown in Fig. 1(b).Table I. Chemical composition of the MAR-M247 powder (wt. %)MARM- 247Cr Co Mo W Al Ti Ta Hf8.0 10.0 0.7 10.0 6.0 1.0 3.0 1.0C 0.015 0.15 0.05BZr NiBal50 m(a)(b)Figure 1. (a) SEM image of the MARM-247 powder, and (b) optical microscopy of the MAR- M247 powder cross-section after polishing.The SLE process was conducted on rectangular EQ MAR-M247 substrates having dimensions of 31.50 mm x 5.84 mm x 2.54 mm. Each substrate was placed into a 31.50 mm x 5.84 mm recess cut into an IN625 base plate. The MAR-M247 powder was placed above the substrates using rectangular wells cut into an Aluminum mask plate.

Each substrate was placed into a 31.50 mm x 5.84 mm recess cut into an IN625 base plate. The MAR-M247 powder was placed above the substrates using rectangular wells cut into an Aluminum mask plate. Once the samples were prepared, they are placed into an atmospheric process chamber that was purged with high purity (99.999%) Argon. A 1kW Ytterbium fiber laser (IPG Photonics, Model: YLS-1000) was used with a Cambridge Technologies galvanometer scanner to focus the beam on top of the substrate to a Gaussian beam diameter of 40 m. A raster scan pattern across the width of the sample generated462a melt pool that linearly propagated along the length of the substrate. All the samples were run with 25.4 m scan spacing.In order to relieve the residual stresses and enable precipitation of the strengthening phases, the samples were heated and kept at 1975 0F for an hour. Thereafter, the samples were cooled at 10 0F /minute or faster to a temperature below 1600 0F.

Thereafter, the samples were cooled at 10 0F /minute or faster to a temperature below 1600 0F. Samples were then given a typical commercial aging treatment consisting of 12 hours at 1600 0F, followed by air cooling at a minimum average cooling rate of 10 0F / minute to 1200 0F to fully develop the strengthening phases.A Buehler automated saw was used to section the samples for microstructural investigation. Each section was mounted in Bakelite and polished to a mirror finish; starting with 80 grit paper and progressively increasing the size to 1200 grit. The samples were then rough-polished using 5 µm and 3 µm diamond solutions. Finally, the samples were polished using a 0.5 µm colloidal alumina suspension. The polished samples were then etched with Marble’s reagent (50 ml HCl, 50 ml H2O, and 10.0 gm CuSO4) to eliminate the γ' phases revealing the dendritic microstructure. Imaging was then completed using a Leica DM6000 optical microscope.

Imaging was then completed using a Leica DM6000 optical microscope. The microstructural investigation of the SLE processed MAR-M247 was carried out on a Hitachi SU8230 SEM.Results and Discussions(a)(b)Figure 2. (a) SEM image showing morphologies of various constituents of SLE processed MARM-M247 deposit in the as-deposited sample and (b) SEM-EDS profile of a typical blocky carbide precipitate in the deposit region of the as-deposited sample.The role of carbides is very complex in nickel-base superalloys. The carbides influence the mechanical properties depending on their morphology and distribution. Fine blocky randomly dispersed carbide precipitates can have strengthening effects on the grain boundaries by inhibiting grain boundary sliding, thus improving creep and rupture strength [12]. On the other hand, if continuous carbide films are formed at the grain boundaries, they might have detrimental effects463on the ductility [12].

On the other hand, if continuous carbide films are formed at the grain boundaries, they might have detrimental effects463on the ductility [12]. Blocky or MC carbides were formed during solidification as discrete particles in the interdendritic regions of the SLE fabricated MAR-M247 as shown in Fig. 2(a). The blocky carbides are very stable at low temperatures but tend to degenerate into secondary carbides at higher temperatures. Figure 2(b) shows the elemental distribution in a typical blocky carbide particle. The carbides were found to be rich in Ta, W, Ti, Hf, and C; however trace amounts of Ni, Cr, Mo, Co, Al, and Cr were also seen to be present. Contrary to the other researchers who reported the presence of Al2O3 [13] in the precipitates, in the current study no such aluminum oxide formation is observed primarily due to the trace amount of oxygen present in the processing chamber.(b)(c)(a)100 mDepositSubstrateElongated CarbideFigure 3.

Morphology of the carbide precipitates in the (a) substrate region, and (b) interface region, and (c) deposit region.Figs. 3(a), 3(b) and 3(c) illustrate the morphology of typical carbide precipitates in the substrate, the interface and the deposit region of the as-deposited sample. The substrate region was abundant with Chinese-script, blocky and elongated carbide precipitates, however in the deposit region, blocky carbides were predominantly observed. However, near the interface, a few elongated carbides were found as shown in Fig. 3(b). This might be due to the fact that during laser processing, a thin layer of the substrate material containing a portion of Chinese-script carbide was464melted. During combined solidification of the partially-melted substrate and the fully melted powder, the existing carbide particles were trapped.

During combined solidification of the partially-melted substrate and the fully melted powder, the existing carbide particles were trapped. Another possible explanation could be convection of the broken Chinese-script carbide arms during laser processing and subsequent solidification of those in the deposit region. The size of the carbide precipitates in the substrate was of order 1-100 m, while in the deposit the size of the blocky carbides was of order 200-400 nm. Chinese-script carbides typically contain B as reported in the literature [14]. However, B was not observed in any of the blocky carbides that were formed in the deposit region. To verify the hypothesis that the elongated carbides found near the interface were actually from the broken Chinese-script carbides from the substrate region, EDS elemental scan was again performed on the elongated carbides. Figure 4 illustrates the elemental maps for the carbide precipitates near the interface.

Figure 4 illustrates the elemental maps for the carbide precipitates near the interface. As can be seen, the elongated carbide precipitate contained only B indicating that the particle was convected from the substrate region.TiTaHfMoCBAlNiCrCoFigure. 4 SEM-EDS elemental maps for various elements in carbide particles deposited near the interface.Figure 5 reports EDS map of various carbide precipitates in the deposit region of the heat- treated sample. The blocky carbides as mentioned earlier were found to be rich in Ta, W, Ti, Hf, and C; however trace amounts of Mo, Zr, and B were also present. In addition to the blocky or MC type carbides, M23C6 carbides forming discrete thin networks also formed in the heat-treated MAR-M247 samples. M23C6 carbides play a crucial role in the strengthening the alloy by inhibiting the dislocation movements. They usually precipitate at the grain boundaries in Cr-containing alloys465as irregular and discontinuous particles.

They usually precipitate at the grain boundaries in Cr-containing alloys465as irregular and discontinuous particles. In the heat-treated MAR-M247 samples, the deposit region showed plate-like carbides that were Cr-rich. Long exposure times at high temperatures resulted in the formation of continuous carbide films along the grain boundaries. Such continuous film formation was shown to affect negatively the ductility and rupture life of the material [12]. In the SLE fabricated MAR-M247, after the heat treatment, only irregular and discontinuous Cr-rich plate-like precipitates were observed.HfTaMoCCrSEMPlate-like CarbideWBlocky CarbideTiBFigure 5. SEM-EDS profile of carbides in the deposit region of a heat-treated sample.ConclusionsIn the present study, SLE deposited MAR-M247 was investigated to characterize carbide formation in the deposited region. MAR-M247 contains about 0.15 wt. % C and 0.015 wt. % B.

MAR-M247 contains about 0.15 wt. % C and 0.015 wt. % B. Minor additions of C resulted in the formation of Ta-rich, MC-type interdendritic primary carbides of predominantly blocky morphologies in the SLE deposited MAR-M247. The blocky carbides were of order 500-1000 nm. Elongated carbides were observed near the interface region. These carbides might be broken pieces of the Chinese-script carbides originated in the substrate region and convected later during combined solidification of the fully melted powder and the partially466melted substrate. The Chinese-script carbides were abundant in the substrate region, however, no such carbide formation was observed in the deposit region. Commercially available heat treatment was carried out, and heat-treated MAR-M247 showed Cr-segregation in the form of a thin plate- like discrete film at the grain boundaries.

Commercially available heat treatment was carried out, and heat-treated MAR-M247 showed Cr-segregation in the form of a thin plate- like discrete film at the grain boundaries. Further investigation will be carried out in the future to characterize the effect of Cr-rich carbide formation on the ductility of SLE fabricated MAR-M247.AcknowledgmentsThis work is sponsored by the Office of Naval Research through grant N00014-14-1-0658.DisclosuresDr. Suman Das is a cofounder of DDM Systems, a start-up company commercializing SLE technology. Dr. Das and Georgia Tech are entitled to royalties derived from DDM Systems’ sale of products related to the research described in this paper. This study could affect their personal financial status. The terms of this arrangement have been reviewed and approved by Georgia Tech in accordance with its conflict of interest policies.References[1] Pollock TM, Tin S.

The terms of this arrangement have been reviewed and approved by Georgia Tech in accordance with its conflict of interest policies.References[1] Pollock TM, Tin S. Nickel-based superalloys for advanced turbine engines: chemistry, microstructure and properties. Journal of Propulsion and Power 2006;22:361-74. [2] Lippold JC, Kiser SD, DuPont JN. Welding metallurgy and weldability of nickel-base alloys: John Wiley & Sons; 2011. [3] Acharya R, Bansal R, Gambone JJ, Das S. A Coupled Thermal, Fluid Flow, and Solidification Model for the Processing of Single-Crystal Alloy CMSX-4 Through Scanning Laser Epitaxy for Turbine Engine Hot-Section Component Repair (Part I). Metallurgical and Materials Transactions B 2014;45:2247-61. [4] Acharya R, Bansal R, Gambone JJ, Das S. A Microstructure Evolution Model for the Processing of Single-Crystal Alloy CMSX-4 Through Scanning Laser Epitaxy for Turbine Engine Hot-Section Component Repair (Part II).

[11] Zeisler-Mashl K, Pletka B. Segregation during solidification in the MAR-M 247 system(of nickel-base superalloys). 7th International Symposium on Superalloys 1992:175-84. [12] Sims C, Stoloff N, Hagel WC. Superalloys II: High Temperature Materials for Aerospace and Industrial Power. 1987. NY: John Wiley & Sons. [13] Bi G, Sun C-N, Chen H-c, Ng FL, Ma CCK. Microstructure and tensile properties of superalloy IN100 fabricated by micro-laser aided additive manufacturing. Materials & Design 2014;60:401-8. [14] Basak A, Das S. A Study on the Effects of Substrate Crystallographic Orientation on Microstructural Characteristics of René N5 Processed through Scanning Laser Epitaxy. 13th International Symposium on Superalloys 2016.468.

rbon nanotube reinforced Polyamide 12 nanocomposites for laser sinteringJiaming Bai*, Ruth D Goodridge†, Richard JM Hague†, Mo SongWolfson School of Mechanical and Manufacturing Engineering, Loughborough University, Loughborough, UK, LE11 3TU †Mechanical, Materials and Manufacturing Engineering, University of Nottingham, Nottingham, UK, NG7 2RD Department of Materials, Loughborough University, Loughborough, UK, LE11 3TU AbstractIn this work, Polyamide12 (PA12) and Carbon nanotube (CNT) added PA12-CNT nanocomposites were laser sintered and investigated. The powder morphology and CNT dispersion of the PA12-CNT were examined. Laser sintering process parameters: powder bed temperature and laser power were studied and optimised. The effect of the addition of CNT on the thermal properties of PA12 was identified.

Laser sintering process parameters: powder bed temperature and laser power were studied and optimised. The effect of the addition of CNT on the thermal properties of PA12 was identified. Compared to the laser sintered parts produced from commercially available laser sintering PA12 powder, the laser sintered PA12-CNT parts showed increased tensile modulus and tensile strength.Keywords: Laser sintering; Polymer nanocomposites; Processing parameters; Thermal conductivityIntroductionOne of the most well established additive manufacturing technology, laser sintering (LS), is a powder-based process which uses a laser as a heat source to fuse polymer powders into three dimensions parts 1. Polyamide 11 and 12 (PA11/PA12) are the most common polymer materials for laser sintering.

Polyamide 11 and 12 (PA11/PA12) are the most common polymer materials for laser sintering. Other polymers such as polypropylene (PP), polystyrene (PS), polycarbonate (PC), poly(ether-ether-ketone) (PEEK) and couple of thermoplastic elastomers are also available from major materials suppliers2,3, however these polymers have not been used widely. Compared to traditional polymer manufacturing processes, such as injection moulding or blow moulding, the materials which are available for laser sintering are limited4. Furthermore, current laser sintering materials can not completely meet the needs, such as mechanical, thermal or electrical requirements, of all products5,6.

Furthermore, current laser sintering materials can not completely meet the needs, such as mechanical, thermal or electrical requirements, of all products5,6. To further develop laser sintering technology, there is an urgent need on researching and developing materials which can meet the various requirements of different applications (automobile, aircraft, aerospace, sports, medical, etc.)Nanofiller-reinforced polymers offer the potential to improve the base material’s properties while remaining processible by conventional processing techniques, and this potential is also expected for the laser sintering process. Several attempts have been made to improve the mechanical or physical properties of base polymers used for laser sintering by adding a nanofiller7,8,9,10,11.

Several attempts have been made to improve the mechanical or physical properties of base polymers used for laser sintering by adding a nanofiller7,8,9,10,11. A common focal point raised in these studies is how to prepare polymer98nanocomposites powders which are suitable for laser sintering4,11.In this study, carbon nanotubes (CNT), allotropes of carbon with a cylindrical nanostructure, were used as the nanofiller due to its outstanding mechanical and thermal properties12. Well-dispersed, laser sinterable PA12-CNT nanocomposite powders with near-spherical particle morphology were fabricated. The melting, crystallisation, and thermal conduction behaviours of both PA12 and PA12-CNT powders were analysed. The dimensional accuracy and mechanical properties of laser sintered parts were also examined.Experimental MaterialsPA12 powder (trade name as ‘PA2200’) was obtained from EOS GmbH, which has an established use in the laser sintering process.

Multi-walled carbon nanotubes were supplied by NanoAmor Materials Inc. A novel method developed in the Materials Department at Loughborough University was used to fabricate near-spherical, well dispersed PA12-CNT nanocomposite powders, which had a CNT content of 0.1% by weight. Thermal conductivity measurementLaser sintering is a heat transfer process, and the ability of the powder material to transfer heat is an important factor for this process. To identify the influence of the CNT on the thermal conduction of the polymer, thermal conductivity of the PA12 and PA12-CNT powders were measured by a thermal conductivity apparatus P5697 (Cussons Technology Ltd, UK).

The apparatus consists of a vertical stack of specimens clamped between an electrically heated source at the top and a water cooled base, all located within a Dewar vessel and furnished with a radiation shield and anti-convection baffle13, shown in Figure 1.The sample specimen was designed as a hollow cylinder (height = 38 mm, diameter = 20mm, wall thickness = 0.5 mm) and produced by laser sintering, in which the testing powders were stored. Two small thermocouples were inserted into the specimen cylinder to contact with the testing powders inside. The inlet and outlet water temperatures were recorded. The heating current was supplied from a variable voltage power pack, which was used to vary the testing temperature. In this work, testing temperature was chosen from 100 – 175oC, with an interval of 25oC.99Heat sourceReferenceT4T3SpecimenT2T1Figure 1: The internal structure of the Dewar vessel.

In this work, testing temperature was chosen from 100 – 175oC, with an interval of 25oC.99Heat sourceReferenceT4T3SpecimenT2T1Figure 1: The internal structure of the Dewar vessel. T1 – T4 are the thermocouples.The thermal conductivity is calculated as:𝐾 =𝐶𝑃 × 𝑀 × 𝐿 × (𝑊2 − 𝑊1) 𝐴 × 𝑡 × (𝑇4 − 𝑇3)where 𝐶𝑃 is the specific heat capacity of water (4186 Joules/Kg), 𝑀 is the mass of water collected in a certain time (t), 𝐿 is the distance between the two thermocouples (T1 and T2), 𝑊1 is the water inlet temperature, 𝑊2 is the water outlet temperature, 𝐴 is the cross-section area of the specimen, 𝑡 is the time for collection of 𝑀, 𝑇1 to 𝑇4 are the thermocouple temperatures13. Laser sinteringBoth PA12 and PA12-CNT nanocomposites powders were laser sintered on an EOS P100 Formiga system.

In terms of ease of processing, different laser sintering processing parameters, including powder bed temperature and laser power were investigated, which are discussed in details in the Results and Discussion section.Tensile test specimens for PA12 and PA12-CNT were fabricated to establish the mechanical properties in accordance with ASTM D638-9914. All test specimens were oriented with the longest dimension aligned vertically to the direction of the movement of the recoating blade, as shown in Figure 2.100Recoat bladeTensile partsyxFigure 2: Schematic of the powder spreading direction in relation to the orientation of the parts Other characterisationsScanning electron microscopy (LEO440 SEM, Leo Electron Microscopy Ltd) was used to image the morphologies of the PA12 and PA12-CNT powders and the dispersion of the CNT in the polymer matrix. Differential scanning calorimetry (SHIMADZU DSC-60) was used to study the thermal behaviours of these powders.

Differential scanning calorimetry (SHIMADZU DSC-60) was used to study the thermal behaviours of these powders. The tensile properties of laser sintered parts were measured on a Zwick 103 testing machine.Results and Discussion Particle morphologyFigure 3 shows the SEM micrographs for the PA12 and PA12-CNT nanocomposite powders. It can be seen that the commercial PA12 powder (Figure 3a) had a near-spherical morphology. The PA12-CNT powder (Figure 3b) maintained the same regular, near-spherical morphology as PA12. Powders with near-spherical and regular morphologies tend to arrange themselves more efficiently, which increases the density of parts during laser sintering process4.

Powders with near-spherical and regular morphologies tend to arrange themselves more efficiently, which increases the density of parts during laser sintering process4. Furthermore, near-spherical powders can flow well and spread uniformly on the powder bed, which is important for the continuous laser sintering process where new powder is deposited each layer.101(a)PA12(b)PA12-CNT20 m20 mFigure 3: SEM micrographs of PA12 powder (a) and PA12-CNT nanocomposite powder (b) Dispersion of CNTFigure 4 shows the FEG-SEM image of the PA12-CNT nanocomposite powders, in which the CNT (white strands) were dispersed uniformly in the PA12 matrix, and no obvious agglomerates were noticed. It is known that nanofiller has a nature to agglomerate together, which could reduce the reinforce effect on the polymer matrix, or even weak the properties of the base polymer15.

It is known that nanofiller has a nature to agglomerate together, which could reduce the reinforce effect on the polymer matrix, or even weak the properties of the base polymer15. Therefore, when nanofiller is used to reinforce polymer matrix, achieving well dispersed, agglomerates free nanofiller in the polymer matrix is very important16,17. Figure 3b and 4 indicated that a near-spherical, well dispersed CNT filled PA12 nanocomposite powders were fabricated successfully.Figure 4: SEM micrograph of CNT (white strands) dispersion in the PA12 matrix Laser sintering process parameter optimisationIn the laser sintering process, the powder bed temperature is controlled to prevent ‘curling’ or ‘caking’. Curling occurs when the powder bed temperature is too low, which causes a high thermal gradient between the sintered and un-sintered material.

Curling occurs when the powder bed temperature is too low, which causes a high thermal gradient between the sintered and un-sintered material. Caking happens when the powder bed temperature is too high, and the powder particle surface starts to melt, which leads to powder agglomeration18. When there is no curling and caking, the powder bed102temperature should be kept as high as possible to achieve optimum mechanical properties with relatively low laser input. In practice, selecting the precise powder bed temperature for laser sintering is highly dependent on the choice of polymer due to the complex thermal properties of different materials. To determine the optimum powder bed temperature for both PA12 and PA12-CNT, it was initially set to 150oC and increased in 10oC increments until the powder bed became hard or agglomerate lumps formed. When this occurred the temperature was reduced 1oC at a time until free powder flow and a smooth, soft powder bed was obtained again.

When this occurred the temperature was reduced 1oC at a time until free powder flow and a smooth, soft powder bed was obtained again. By this method, the suitable powder bed temperature for PA12 and PA12-CNT were determined as 171oC and 172oC respectively.In addition to powder bed temperature, laser power is another crucial parameter. Different laser powers were investigated, with the laser scan speed and laser scan spacing being maintained at 2500 mm/s and 0.25 mm respectively. Initial attempts to build parts using relatively low laser powers, 13W and 15W, produced parts whose edges curled up in the initial layers, which lead to failure of the build as the recoater could not pass over the raised section. When the laser power was increased to 17W, there was considerably less curl. When the laser power was increased further to 21W, good definition and no distortion parts were built successfully for both PA12 and PA12-CNT.

When the laser power was increased further to 21W, good definition and no distortion parts were built successfully for both PA12 and PA12-CNT. The summary of the processing parameters used in this study are shown in Table 1.Laser sintering parameterPA12PA12-CNTPowder bed temperature (oC)171172Layer thickness (mm)0.10.1Laser power (W)2121Laser scan speed (mm/s)25002500Laser scan spacing (mm)0.250.25Table 1: Laser sintering parameters for PA12 and PA12-CNT powdersThe laser sintering of tensile specimens is shown in Figure 5. The specimens for both PA12 and PA12-CNT were built successfully with good definition.

The specimens for both PA12 and PA12-CNT were built successfully with good definition. Apart from the colour difference, there were no obvious differences between the PA12 and PA12-CNT laser sintered parts visible by eyes.103(a)(b)Figure 5: Laser sintering of (a) PA12, and (b) PA12-CNT nanocomposites Thermal analysis of powdersThe melting and cooling characteristics from the DSC curves for the PA12 and PA12-CNT powders are shown in Table 2. There was very small difference (0.6oC) on the melting point between the PA12 and PA12-CNT. The crystallisation temperature of the PA12-CNT nanocomposites powder was 5oC higher than the PA12, which suggested that the CNT had a strong nucleation effect on the PA12. Laser sintering processing window, which is the temperature range between melting and crystallisation temperature, should be wider enough to ensure the feasibility of laser sintering.

Laser sintering processing window, which is the temperature range between melting and crystallisation temperature, should be wider enough to ensure the feasibility of laser sintering. The processing window for the PA12-CNT was decreased about 5.6oC compared to PA12; however it is still wide enough, which was confirmed during the laser sintering process in the previous section. The crystallinity of both PA12 and PA12-CNT powders were also measured. A theoretical enthalpy of melting for 100% crystalline PA12, ΔHC100% = 209.2 J/g was used for calculation19.Sample Melting point (oC)Crystallisation Temperature (oC)Crystallinity (%)PA12183.4146.627.4PA12-CNT184.0151.626.5Table 2: Melting, crystallization and crystallinity of PA12 and PA12-CNT powdersThe results of the thermal conductivity of the PA12 and PA12-CNT powders are shown in Figure 6. It can be seen that as the temperature rise, the thermal conductivity of both PA12 and PA12-CNT increased slightly.

It can be seen that as the temperature rise, the thermal conductivity of both PA12 and PA12-CNT increased slightly. PA12-CNT powders appeared average 14.2% greater thermal conductivity than PA12. This might be explained by the remarkable higher thermal conductivity of CNT, which could be over 3000 W/mK20, compared to PA12. Increased thermal conductivity can improve the laser absorption, which could lead a more efficient powder fusion and melting process.104) K m / W( y t i v i t c u d n o cla m r e h T0.18 0.17 0.16 0.15 0.14 0.13 0.12 0.11 0.1 0.09 0.08 0.07 0.06 0.05PA12 PA12-CNT75100125 Temperature (oC)150175200Figure 6: Thermal conductivity of PA12 and PA12-CNT powders Dimensional accuracy of laser sintered partsShrinkage is a common phenomenon for laser sintered parts, which can reduce the dimension accuracy.

The dimensional accuracy of PA12 and PA12-CNT laser sintered parts was investigated by comparing predefined part dimensions and laser sintered part dimensions. Both parts showed shrinkage after laser sintering, which caused the dimensional error (Table 3). Compared to PA12, PA12-CNT laser sintered parts showed better dimensional accuracy.MaterialsDimensional error (%)PA127.9PA12-CNT3.7Table 3: Dimensional error of laser sintered PA12 and PA12-CNT parts Mechanical characterisation of laser sintered partsThe tensile properties of the PA12 and PA12-CNT laser sintered parts are listed in Table 4. PA12-CNT had 44.5% increased tensile modulus and 7.0% improved tensile strength compared to PA12.

PA12-CNT had 44.5% increased tensile modulus and 7.0% improved tensile strength compared to PA12. This enhancement may come from two aspects: one is the load transfer between PA12 matrix and the CNT, which indicated strong polymer-filler interaction; the other is the enhanced laser absorption of PA12-CNT, which was due to the higher laser absorption coefﬁcient and thermal conductivity of CNT.105Tensile propertiesPA12PA12-CNTTensile modulus (MPa)2554.85 ± 78.593691.85 ± 92.51Tensile strength (MPa)49.96 ± 1.3753.45 ± 1.25Table 4: Tensile properties of laser sintered PA12 and PA12-CNT partsConclusionA CNT-filled PA12 nanocomposite was prepared and laser sintered successfully. The PA12-CNT powder had near-spherical morphology with well-dispersed CNT in the polymer matrix.

The PA12-CNT powder had near-spherical morphology with well-dispersed CNT in the polymer matrix. The thermal analysis showed that the thermal conductivity increased slightly with temperature increase for both PA12 and PA12-CNT powders; the PA12-CNT powders had greater thermal conductivity than PA12. Laser sintered PA12-CNT parts showed less dimensional error than PA12 parts.

Laser sintered PA12-CNT parts showed less dimensional error than PA12 parts. Tensile results showed 44.5% greater tensile modulus and 7.0% higher tensile strength for PA12-CNT, which might be explained by the load transfer between the polymer and nanofillers, and the enhanced laser absorption.References1 Hopkinson N, Hague RJM, Dickens PM, Rapid Manufacturing: An industrial revolution for the digital age, England: John Wiley & Sons Ltd, 20052 http://www.eos.info/en/products/materials/materials-for-plastic-systems.html, accessed on July 2012.3 http://production3dprinters.com/materials/sls, accessed on July 2012.4 Goodridge RD, Tuck CJ, Hague RJM, Laser sintering of polyamides and other polymers, Progress in Materials science, 2011, 57(2) 229 – 267.5 Kruth JP, Levy G, Klocke F, Childs T, Consolidation phenomena in laser and powder-bed based layer manufacturing.

viewed PaperA CASE STUDY IN COMPONENT REDESIGN FOR ADDITIVE MANUFACTURING PROCESS WORKFLOWSElliott Jost*, Jaime Berez*, Christopher Saldaña*George W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology, 801 Ferst Drive NW, Atlanta, GA 30332AbstractAdditive manufacturing (AM) has matured beyond limited use-cases in rapid prototyping into a process capable of competing with conventional manufacturing methods in the production of end-use components. As such, many manufacturers are evaluating candidate products for redesign for AM, with interest in improving component performance, streamlining manufacturing, and reducing costs. In this study, the authors argue for systematic use of opportunistic design tools such as generative design and for the inclusion of restrictions imposed by the entire manufacturing process workflow to be included in the design process.

Emphasis is placed on how secondary processing steps, that is, those other than the primary AM process, inform design. A bicycle stem is used as an exemplar case study of component redesign for AM. Generative design is used to optimize the component’s weight given the design constraints. An end-to-end manufacturing process chain is consequently developed and analyzed for viability, covering design for additive manufacturing (DFAM) and post processing. Through this comprehensive case study, it is shown that significant weight savings, greater than 25% in the present case, can be achieved through the DFAM process. Guidelines from the DFAM process are generalized for application to further cases.IntroductionAdditive manufacturing (AM) has shifted the manufacturing and design paradigm in recent years, particularly as some AM technologies have matured from rapid prototyping methods to suitability for production-scale manufacturing.

Laser powder bed fusion (LPBF) is an AM technology which can process metals by using a laser heat source to fuse selective areas in a thin bed of metal powder feedstock, repeated in a layer-wise manner, progressively building the workpiece. Of the currently available metal AM processes, LPBF has been rapidly commercialized and has been adopted within industries such as aerospace to medical which utilize LPBF for near- net-shape manufacturing of relative dense geometrically complex components. Additive manufacturing favored in large part because of the significant design freedom it offers, as has been reported [1]. Along with this newfound design freedom comes a significant shift in how designers conceptualize designs, think about redesign of components and assemblies, and utilize tools for design improvement or optimization.

In the present work, the authors argue that the successful design for AM (DFAM) workflow incorporates considerations spanning the entire manufacturing process chain, and that post-processing manufacturing technologies are a crucial, yet commonly neglected aspect of the DFAM process. In the ensuing, a case study is presented in which a bicycle stem formerly manufactured via conventional means is redesigned for LPBF AM. Evaluation of design tradeoffs relating to the entire manufacturing process chain are considered and discussed.1981Design for additive manufacturing (DFAM) has been an object of study since AM technologies were recognized to enable “free complexity” in design [1–4]. Laverne et al. [5] divided DFAM into two categories for design assessment: opportunistic and restrictive design, which will be discussed and applied in the context of the present study.

Laverne et al. [5] divided DFAM into two categories for design assessment: opportunistic and restrictive design, which will be discussed and applied in the context of the present study. Opportunistic design is a design methodology which emphasizes utilization of the geometric and material complexity offered by AM to improve designed components and manufacturing workflows. This includes but is not limited to advanced design and optimization technologies such as topology optimization (TO) and generative design (GD), each of which utilizes computational algorithms to provide improved designs. Conversely, the restrictive DFAM methodology focuses on using the limits of AM such as natural constraints relating to minimum manufacturable feature size, surface finish, or the manufacture of overhanging structures to guide the design workflow. Meisel et al. [6] argued that a successful design utilizes both opportunistic and restrictive design, which is referred to as dual design.

Meisel et al. [6] argued that a successful design utilizes both opportunistic and restrictive design, which is referred to as dual design. Similarly, Hällgren et al. [7] divided the DFAM into process and designer driven, incorporating software tools and human experience into the design process. In process driven DFAM, a focus is put on manual interaction to reduce design time. This can incorporate complex technologies such as TO and GD, as mentioned above. Designer driven DFAM utilizes more simplistic, less deterministic methods. For example, a designer may use lattices to reduce mass on an object as opposed to a more complex topology optimization implementation.Opportunistic, computer-aided design tools such as TO or GD are being increasingly relied upon because of their ability to maximize the potential of AM [8]. Despite this, these software tools are not yet well integrated into design and manufacturing workflows [9].

Despite this, these software tools are not yet well integrated into design and manufacturing workflows [9]. TO and GD provide systematic, rigorous, and mathematical exploring of large design spaces. Topology optimization has been widely studied [10–13] for AM, and utilizes mathematical optimization to converge on optimal designs given an objective function, typically defined as a displacement, structural load, or thermal performance, and a complimentary set of boundary conditions. Similarly, GD, a relatively new design tool enabled by computing power increases in recent years, uses artificial intelligence to create suggested designs that optimize for a specific objective. These tools are useful because, as Wu et al. [14] writes, “even for the most experienced designers, their intuition might be limited when manually exploring” a design space.

These tools are useful because, as Wu et al. [14] writes, “even for the most experienced designers, their intuition might be limited when manually exploring” a design space. Despite the use of both TO and GD in exploring large design spaces, these tools must be used responsibly, and do not currently incorporate all necessary design considerations, as will be demonstrated in the present paper.Despite claims that AM can create fully functional end-use components [6,15], the authors argue that this is the exception and not the rule. Although AM processes, and specifically LPBF, can create end-use components in their as-printed condition in certain scenarios, it is much more common to require significant post-processing. Peng et al. [16] similarly argued that AM faces critical issues inhibiting the direct use of as-printed parts.

Peng et al. [16] similarly argued that AM faces critical issues inhibiting the direct use of as-printed parts. Example scenarios which require significant post-processing include fatigue applications [17–19] and interface surfaces, each of which is well documented in the literature. Manufacturing methods used in these scenarios may include heat treat [20,21], hot isostatic pressing (HIP) [22–24], and surface finishing [25–28]. Consequently, the authors posed that successful DFAM incorporates all aspects of the manufacturing process chain and is not limited to considerations related only to the AM process to be used. This conclusion is consistent with multiple reviews covering topics within DFAM that similarly agree that manufacturing considerations should be further incorporated into the DFAM workflow [5,29–31], and is demonstrated explicitly through the case study redesign in the present1982work.

Furthermore, restrictive design should consider constraints on manufacturing freedom imposed by AM, but also by the subsequent manufacturing technologies used for post-processing necessary to create an end-use component. In the present study, a case example part - a bicycle stem component - is redesigned using both opportunistic and restrictive DFAM strategies with emphasis on the post-processing strategies necessary to fully implement such a component.Redesign approachIn this work, the application of a dual design approach to DFAM is utilized, with an emphasis on how manufacturing processes beyond the AM process supply restrictive input to the design process. That is to say, incorporating manufacturing process planning considerations into design stages (embodiment and detail design) is vital. As is well established by the literature, opportunistic DFAM approaches such as GD should be incorporated in embodiment design to reap the benefits which AM can provide [5,6].

As is well established by the literature, opportunistic DFAM approaches such as GD should be incorporated in embodiment design to reap the benefits which AM can provide [5,6]. Additionally, heuristics for DFAM need to be applied as restrictions to ensure manufacturability. Here, it is emphasized that AM processes for metals such as LPBF are rarely the sole manufacturing process required to produce a functional product. The LPBF process is well understood to produce geometry with poor dimensional control [32], gross deformation due to residual stress, high surface roughness [33], low microstructural heterogeneity [34], and high volumetric defect content [35]. As such, secondary manufacturing processes are often leveraged to address these shortcomings [16]. Further, the sequence of secondary processing steps suitable to an AM component may be product- or application-specific and does not necessarily mirror classical approaches to manufacturing process planning.

The design process must incorporate this knowledge from an early stage for the DFAM approach to supply viable and competitive alternatives to conventional manufacturing workflows.Figure 1 shows the proposed workflow for DFAM. At top, the overall design process is coarsely shown, which includes problem definition, conceptual design, embodiment design, and detail design [5,36]. The additional step of manufacturing process planning has been added, which is connected by arrows to various stages of the design process to emphasize the iterative nature of the design process and post-processing considerations necessary for consideration in the design. Note that while this step is shown to be strictly outside of the design process as classically defined, this work argues for its close relation. Several of these high-level steps are enclosed within a dotted box, and these steps are expanded upon in detail in the middle portion of the figure.

Several of these high-level steps are enclosed within a dotted box, and these steps are expanded upon in detail in the middle portion of the figure. Since the case study presented is for a relatively simple component-level redesign, the conceptual design step is not emphasized. Rather, embodiment and detail design are focused on. These stages are expanded upon in the chart, and are comprised of configuration & parametric design, DFAM, simulation & down selection, and design for manufacturing (DFM). DFAM is shown as informing configuration and parametric design (note the left-pointing arrows) based on dual inputs: opportunistic and restrictive. Opportunistic inputs might support the selection of a GD design with organic and complex geometries, for example. Restrictive inputs will inform material selection, the minimal feature size, or the design of self-supporting geometry, for example.

Restrictive inputs will inform material selection, the minimal feature size, or the design of self-supporting geometry, for example. The incorporation of these DFAM principles is iterative, as illustrated below by the loop formed in the flow chart. Simulation and down selection follows as a highly flexible step which may vary on a case-by-case basis. Simulation may be performed to validate a GD design, test the robustness of a design against a variety of parametrically changed loading scenarios, or test performance in the presence of small design modifications. Down selection between multiple GD designs or lattice variations may also1983be required. DFM is a subsequent step which informs all prior steps. Here, designers must acknowledge in detail the likely necessity of secondary processing steps and incorporate these considerations iteratively. Finally, at the lower portion of the figure, a generalized manufacturing process workflow where AM is the primary process, is shown.

Finally, at the lower portion of the figure, a generalized manufacturing process workflow where AM is the primary process, is shown. The intimate and detailed knowledge of the manufacturing process workflow undergirds the DFAM and DFM steps, shown via the dotted arrows.Figure 1. Generalized AM design and manufacturing process planning workflow.In the lower portion of Figure 1, which encompasses manufacturing process planning, the primary manufacturing process, AM, feeds into an unordered set of potential secondary processes. Some of these processes are semi-specialized to AM or even the selected AM process, such as build plate separation or HIP. Others are more common secondary steps in manufacturing workflows, such as machining, heat treat, surface finishing, or quality and non-destructive evaluation (NDE). Even so, there are still AM-specific considerations to be addressed.

Even so, there are still AM-specific considerations to be addressed. Surface finishing may take on greater importance and require novel approaches due to the characteristically rough surfaces produced by AM [32,33]. Quality and NDE steps such as X-ray computed tomography (XCT) may be essential due to the defect prone nature of AM [34,35,37] and complex geometries difficult to inspect via traditional methods. Stress relief, as a heat treat operation, may be required more frequently for AM parts than is typical. Machining may tend to leverage certain processes such as EDM due to its nature as a zero-force machining method appropriate for delicate geometries and insensitivity to high hardness, low-machinability materials popular for use with AM such as stainless steels, nickel- and cobalt-based super alloys, and refractory metals [38,39]. Finally, the ordering of each of these steps will likely be specific to the material, geometry, AM technology, and design application considered.

Finally, the ordering of each of these steps will likely be specific to the material, geometry, AM technology, and design application considered. This ordering will be further discussed in the subsequently presented case study.1984An exhaustive list of secondary processing concerns for AM products would be out of scope for a single study. In addition to the generalized explanation above, this works presents a case study which illustrates specific examples of how a design process for additively manufactured products requires not only DFAM principles but supporting DFM principles. Critically, these considerations are specific to the integration of AM in a larger manufacturing process plan.Case studyTo better understand and illustrate the DFAM process as discussed above, a bicycle stem was chosen as an exemplar part to be redesigned. A bicycle stem serves as the connection between the steerer tube on the front fork and the handlebars, as shown in Figure 2.

A bicycle stem serves as the connection between the steerer tube on the front fork and the handlebars, as shown in Figure 2. A bicycle stem is intended to be lightweight yet stiff to provide the rider with efficient control of the handlebars. Thus, the objective of the case study is to redesign a traditional bicycle stem for AM with the goal of minimizing weight while still meeting the design requirements.Figure 2. Bicycle stem schematic description.Problem definition: Design constraintsService requirements in the form of loading conditions were determined through a literature search of loads experienced on a bicycle during common riding scenarios [40,40,41]. Three categories of scenarios were identified including starting, braking, and hitting a pothole. Loads corresponding to these scenarios are detailed in Table 1 below.

Three categories of scenarios were identified including starting, braking, and hitting a pothole. Loads corresponding to these scenarios are detailed in Table 1 below. Starting 1 and 2 are mirror images of each other because of the asymmetric loading experienced during the start depending on whether the rider begins with force on the left or right pedal. The loading conditions were adapted to represent those experienced for a 90 kg (200 lbs) rider. The force and moment directions noted correspond to the coordinate system used in Figure 2, where the Z direction is vertical, X is along the forward movement direction of the bicycle, and Y is mutually perpendicular. Note that the stem has a tilt of 5° from the X direction, as seen in Figure 2, which is common practice to improve the ergonomic position of the handlebars.Table 1.

Note that the stem has a tilt of 5° from the X direction, as seen in Figure 2, which is common practice to improve the ergonomic position of the handlebars.Table 1. Bicycle stem loading conditions for 90 kg riderStarting 1 Starting 2 Braking PotholeForce (N) Fx -174 -174 448 500Fy -58 58 0 0Fz 417 417 -172 -1250Moment (N·m) My Mx 0 -211 0 211 0 0 0 0Mz -19 19 0 01985Embodiment designIn-depth conceptual design was not emphasized for this case study since the existing conceptual design for a bicycle stem is well defined. Furthermore, because the present case study is a redesign of an already existing product, this work’s focus has been put on the later stages of the design process. As such, this case study begins with embodiment design.

As such, this case study begins with embodiment design. Embodiment design as defined in Figure 1 includes material selection, early CAD product definition, and GD optimization, each of which are discussed below.Material selectionBecause weight is to be minimized in this redesign, a high specific property material is desirable. In other words, materials with high stiffness or strength on a per weight basis are desired. Early conceptual design identified metal alloys as the candidate material class for this design. Non- structural alloys, e.g., copper alloys, tool steels, precious metals, refractory alloys, were also excluded in early conceptual design. At the embodiment design stage, a process-first approach to material selection was used, providing restrictive DFAM input to the design process.

At the embodiment design stage, a process-first approach to material selection was used, providing restrictive DFAM input to the design process. Materials compatible with the LPBF process were considered, restricted to those which are commercially available in suitable powder form from multiple suppliers, including prevalent ferrous alloys, e.g., stainless steel 17-4 and 316, non-ferrous alloys, e.g., aluminum AlSi10Mg and titanium Ti-6Al- 4V, and super alloys, e.g., Inconel 718. Material selection was performed using material performance indices in the form of Ashby charts. The indices selected were σy/ρ and E/ρ, where σy is yield strength, E is the modulus, and ρ is the material density. Ashby charts which plot modulus and yield strength against density are shown in Figure 3. These indices identify the most mass- efficient material to maximize strength and stiffness for a tensile loading application.

These indices identify the most mass- efficient material to maximize strength and stiffness for a tensile loading application. While the stem is subject to applied forces which create net bending and torsion, the developed GD geometry has strut-like geometrical elements which are primarily loaded in tension. This provides an interesting example of how opportunistic DFAM prompts different material selection approaches than conventional manufacturing. Should a design, such as that shown in Figure 2, be considered, 2/3/ρ, G1/2/ρ, and E1/2/ρ may be more appropriate to identify the material selection indices such as σy most mass-efficient material for torsion and bending [36].Figure 3. Ashby charts used for material selection1986As shown in Figure 3, of the commercially available alloys suitable for the LPBF process, AlSi10Mg and Ti6Al4V (Ti64) have similar specific moduli at 26.6x106 m2/s2 (Pa/[g/cm3]) and 25.7 x106 m2/s2, respectively.

However, Ti64 has by far the highest yield strength per density value of 199x106 m2/s2. Consequently, Ti64 was selected as the material for use in this study.Generative design optimizationGenerative design, a category of tools utilizing artificial intelligence to create optimized designs, has been chosen to augment the human designer’s capabilities. Specifically, Autodesk Fusion360 Generative Design was used in the present redesign to reduce the mass of the bicycle stem. To reduce the computational time, a simplified model of the bicycle stem without fillets, bolt holes, etc. was utilized. This had the additional advantage of creating a more design-agnostic bicycle stem model that did not assume a particular way of fixturing the bicycle stem to the handlebar and steerer tube. For the GD tool, volumetric regions of the design needed to be assigned to one of three categories: preserve, obstacle, or starting, as shown in in Figure 4(a).

For the GD tool, volumetric regions of the design needed to be assigned to one of three categories: preserve, obstacle, or starting, as shown in in Figure 4(a). Preserve geometries are those regions that should not be acted upon by the GD tool such as mating surfaces, connections, etc. Starting geometry provides a reference point for the GD tool to begin designing from and is heavily influential on the converged upon designs, though its effects are not studied here. Obstacle geometry was placed in regions occupied by mating components, namely the handlebars and steerer tube.Figure 4. (a) Design space region categorization additionally showing a generalized case applied loads on the handlebar mount. (b) Side view showing an example of an applied bearing load on the handlebar mount. (c) Front view showing an example of a resolved moment on the handlebar mount.Loads were applied within the GD tool using the built-in software tools.

(c) Front view showing an example of a resolved moment on the handlebar mount.Loads were applied within the GD tool using the built-in software tools. The inner surface of the steerer tube mount was fixed in all degrees of freedom. To simulate the application of loads as realistically as possible, the forces were applied as bearing loads, which have approximately parabolic force distributions, as illustrated in Figure 4. Moments were modified to be applied as forces at the edges of the handlebar mounts due to limitations in the application of moments within the software. Load cases were applied as individual scenarios within the GD tool, however the1987software factors in all scenarios in proposed designs.

Load cases were applied as individual scenarios within the GD tool, however the1987software factors in all scenarios in proposed designs. An example of how various bearing loads or moments would be applied is illustrated in Figure 4(b-c).In addition to design space categorization and load cases, the GD tool required inputs for the design objective, print orientation, maximum overhang angle, material properties, minimum feature thickness, and safety factor. The objective was set to minimize mass. Minimum feature thickness was set to 1.5 mm and maximum overhang angle to 45 degrees based on the known performance of the EOS M280, which was the exemplar commercial LPBF platform selected for this case study. Print orientation was specified as the +X direction as defined Figure 4. The material properties used corresponded to the Ti64 parameters shown in Figure 3.

Print orientation was specified as the +X direction as defined Figure 4. The material properties used corresponded to the Ti64 parameters shown in Figure 3. The safety factor was set to 2.0, where exceeding the yield strength of 881 MPa was the service criterion.Based on the four load cases applied to the model and the requirements described above, the GD software generated several models that met the specified loading and manufacturing criteria. These designs were evaluated based on their weight savings and manufacturability. The chosen design is shown in Figure 5a, and was selected due to the significant weight savings through optimized material location. It should be noted that this intermediate design includes no integration of attachment points, or other general DfAM principles beyond the maximum overhang angle and minimum feature size constraints input to the GD software.

Using the intermediate design shown above, the simplistic handlebar and steerer tube mounts were modified to include the bosses for pinch bolts as in the original design. In later manufacturing steps, both the handlebar and steerer tube mounts are split in a machining operation, thereby allowing the pinch bolts to apply clamping force. This modified design is shown in Figure 5b.Figure 5. (a) Intermediate bicycle stem design as output from the GD software tool. (b) Design as-modified to have bolt holes and attachment points as in the original stem design.Detail designThe detail design phase of the design process includes detailed DFAM specific to the intermediate, embodiment design presented above in Figure 5. Specifically, considerations for limitations of the process, ease of manufacture, and preparation for secondary processing are discussed.1988Design for Additive ManufacturingOnce the GD step was complete, DfAM considerations could be incorporated in further detail.

This was done by evaluating the design shown in Figure 5b and making appropriate modifications to improve manufacturability for quality, speed, and ease. As specified in the generative design tool, the choice was made that the build direction would be parallel to the axial direction of the stem. With this part orientation, 20 build stems will be able to be made on a single 250 x 250 mm build area, the maximum build area of the EOS M280. Table 2 shows several DfAM heuristics that were considered and implemented. Discussion of tradeoffs and further considerations are presented below.Table 2.

Table 2 shows several DfAM heuristics that were considered and implemented. Discussion of tradeoffs and further considerations are presented below.Table 2. DfAM heuristics and resulting solutionsDfAM Heuristic Features should be <1.5 mm for reliable process performance and geometrical control Horizontal holes must be < 8mm diameterParts should be firmly attached to buildplate to sink heat and resist thermal deformation Overhangs must be less than 45 degrees (from build direction) LPBF as-built surface texture not acceptable for mating surfaces Minimize height to reduce build timeSolution/Implementation Enforce through generative design toolAdd teardrop profile to circular bores above this size Volumetric support added to bottom of steerer tube mount Bolt hole boss modification, anchoring support on steerer tube mount Add machining allowances to critical mating surfaces.

Favor more parts per build rather than reduced build time Remove need for supports where possibleThe use of fewer supports reduces post- processing timeDue to constraints input to the GD tool, the selected design meets the minimum feature size requirements. Furthermore, the previously selected build direction orients the struts or ligaments of the design at an acceptable build orientation to not require supports. The output of the GD tool is shown in stage 1 of Figure 6. This design was modified to include additional functional features in a detail design stage, as seen in stage 2 of Figure 6. As already discussed, bosses for pinch bolts were added. Additionally, the clamps were split across these bosses to enable the pinch bolts to properly function. This stage of design represents an early engineering product definition, as it might be used ‘in-service’ on the bicycle.1989Figure 6.

This stage of design represents an early engineering product definition, as it might be used ‘in-service’ on the bicycle.1989Figure 6. DFAM and DFM process.Further evolution of this geometry is required to develop an ‘as-built’ design suitable for subsequent manufacturing steps – seen in stage 3 of Figure 6. As with all LPBF parts, the bicycle stem will need to be anchored to the build plate. For simplicity and ability to sink heat away from the part, a simple projected area solid volume support has been added. This support geometry will also provide allowance for separating the component from the build plate and will be removed in the finish machining process. A tradeoff of the selected build orientation is that the large bores on the handlebar and steerer tube mounts are oriented horizontally, and thus cannot be directly printed.

A tradeoff of the selected build orientation is that the large bores on the handlebar and steerer tube mounts are oriented horizontally, and thus cannot be directly printed. To address this, the bores were modified to have a ‘teardrop’ profile, which prevents any internal surface from having an overhang greater than 45 degrees. A radial machining allowance of 1.25 mm was added to the bores as well so that the surface finish requirements can be met for the mating between the mounts and opposing surfaces. Diametrical allowances of 1.0 mm were added to the bolt holes to allow for drilling and tapping the holes to the proper requirements. Finally, the splits on the clamps were removed. The split on the handlebar clamp (top of the figure) cannot be integrated into the LPBF process, as it produces overhanging and floating geometry. While the steerer tube clamp split (bottom of the figure) could be manufactured with LPBF, it is also removed. This is due to a DFM consideration.

While the steerer tube clamp split (bottom of the figure) could be manufactured with LPBF, it is also removed. This is due to a DFM consideration. The inner surface of the clamp must be finish- machined (discussed in detail later) and the presence of the split would prevent rigid work holding and produce chatter in cutting.After this round of DFAM- informed design, a manufacturability analysis, seen in stage 4 of Figure 6, was performed using the software tool Materialise Magics, which is used to design supports and otherwise prepare files for AM. This analysis showed that the supports required to build the handlebar mount bolt hole bosses would be significant in size, intersect other stem geometry, and be difficult to remove in post-processing. As such, the DFAM process was repeated1990iteratively, resulting in the geometry shown in stage 5 of Figure 6. Modifications were made to the bolt holt bosses, angling them such that supports were not needed.

Modifications were made to the bolt holt bosses, angling them such that supports were not needed. This results in final as-built CAD product definition, previously based on two intermediate models. This product definition has incorporated major restrictive DFAM considerations, after originally being primarily generated based on opportunistic principles. Note the extreme differences between the as-built product definition and the early in-service product definition.Manufacturing process planningTo address all restrictive inputs relevant to the entire manufacturing process workflow, the following section details secondary post-processing steps necessary for the present case study. This final step of the design process is crucial for the creation of a component suited to mass production and functionality in service.

This final step of the design process is crucial for the creation of a component suited to mass production and functionality in service. The specific steps and the ordering of these steps is part and process-specific, as discussed in the introduction.Heat treatment, surface finishing, and build plate separationAfter printing, loose powder removal will be accomplished using a combination of industrial vacuuming and manual cleaning. Due to the lack of enclosed or tight spaces more rigorous methods of powder removal such as ultrasonic cleaning should not be necessary. Following powder removal, heat treatment can commence. Heat treatment is required for the bicycle stem part for two reasons: 1) to remove residual stresses induced by the rapid heating and cooling cycles innate to the LPBF process, and 2) to promote phase transformation and precipitation to produce optimal material strength.

Heat treatment is to be performed on the build plate due to the restraint it offers against residual stresses. Removing the workpiece prior to a stress relief operation will lead to deformation. The prescribed heat treat is a two-part process as follows and was informed by AMS 2801BC. First, the parts are to be solutionized for 15 min at 955C (1750 F), then water quenched. This temperature is well over the static recrystallization temperature for Ti64 and will allow for stress relieve. Quenching will transform much of the high temperature microstructure to an alpha-prime martensitic phase, which will strengthen the alloy. After the part has been quenched, an aging treatment at 482 C (900 F) for 8 hours is prescribed, followed by air cooling. This intermediate temperature tempers unstable martensite and will precipitate fine beta particles, strengthening the material via precipitation hardening.

This intermediate temperature tempers unstable martensite and will precipitate fine beta particles, strengthening the material via precipitation hardening. Lastly, bead blasting is needed to remove scale, partially fused satellite particles, and to provide a more uniform and smoother surface finish on the stems. After heat treat the part is removed from the build plate. Due to the solid volume support connecting the part to the build plate, a low-precision high speed approach is viable. Band sawing will be used, as this is a more cost effective and faster method than wire-EDM.Finish-machiningA bicycle stem has several features that require precise dimensional control and good surface finish not within the capabilities of LPBF. These features include the bore mating surfaces that clamp the stem to the steerer tube and the handlebars.

These features include the bore mating surfaces that clamp the stem to the steerer tube and the handlebars. To clamp these with evenly distributed pressure that does not unduly stress the stem, the internal diameter of the bore must be controlled with tight tolerance. Additionally, high surface roughness can damage the clamped steerer tube1991and handlebars. A machining process was developed for use with a two-station vise on a 3+1 axis or higher degree of freedom vertical machining center. The overall process requires two set-ups. To appropriately hold the bicycle stem, two pairs of ‘soft jaws’ that conform to the as-built outer profile of the handlebar and steerer tube mounts will be manufactured for each of the two work- holding stations. In the first set-up the steerer tube clamp is fixtured in soft jaws so that it is rigidly fixtured, and the handlebar clamp is cantilevered and therefore not rigid.

In the first set-up the steerer tube clamp is fixtured in soft jaws so that it is rigidly fixtured, and the handlebar clamp is cantilevered and therefore not rigid. Thus, the second station reverses the part, fixturing the handlebar mount so that it may be machined. Five distinct operations are shown and summarized in Table 3 for the steerer tube mount side of the stem. Similar operations are performed on the handlebar mount side except for the build plate support removal. In summary, the bulk of machining allowances within the bore are first removed using an adaptive clearing strategy. Then, finish boring is performed with small depth of cuts to create a precise, high surface finish bore. Next the build plate support is machined away using an adaptive clearing strategy and bolt holes are drilled and tapped. Finally, a slitting operation is performed so that the mount may properly clamp the steerer tube or handlebars when the pinch bolts are tightened.

Finally, a slitting operation is performed so that the mount may properly clamp the steerer tube or handlebars when the pinch bolts are tightened. This operation requires a plug for the machined bore to ensure no deformation under fixturing loads while machining.Operation 1 Rough milling of boreTable 3. Machining process plan overview Operation 3 & 4 Removal of support, drilling and tapping of holesOperation 2 Finish boringOperation 5 Slitting5 min, 34 sec6 min, 23 sec1 min, 15 sec6 min, 8 secTechnical evaluationTo evaluate the technical fitness of the design, a static finite element (FE) model was created using Abaqus/CAE to evaluate the stress distribution experienced by the bicycle stem during the four examined loading scenarios. Although the FE model cannot validate the design entirely or itself be validated without experimental testing not performed for this case study, it can still be useful as a secondary check of the generative design tool output.

Acknowledging these limitations, the details of the modeling approach are provided below.Boundary conditions were applied to the bicycle stem model in a similar manner to the GD tool. However, due to slight differences in the GD and FE tools, a reference point in the center of the handlebar mount was defined and rigidly constrained to exist at the center of the bore during the small deformations occurring due to loading. Either a point force or point moment is applied1992to the reference point depending on the loading case, which simulated application of loading conditions detailed in Table 1. All translations and rotations of the steerer tube mount were constrained to fix the model in space and simulate a fixed steerer tube mount.The model was meshed using C3D10 quadratic tetrahedral elements with a mesh size of 0.2 mm.

A mesh convergence study, not shown here for brevity, was performed, and verified that this mesh size is sufficiently fine to capture the bulk stress-strain response of the bicycle stem under the specified loading conditions. Analysis of the four loading scenarios revealed that the R tool produced a design that with stress distributions consistent with design specifications. Figure 7 shows the resulting stress distributions along on the deformed geometry for each loading scenario. Note that deformations are exaggerated for visual effect. As can be seen in this figure, particularly in the “starting” scenarios, which represent the most severe loads observed by the bicycle stem, the stress distribution is uniform throughout the GD-optimized region. This can be seen by the consistent green stress value within this region.Figure 7. Finite element model results showing nominal stress distributions in samples.

This can be seen by the consistent green stress value within this region.Figure 7. Finite element model results showing nominal stress distributions in samples. Deformations exaggerated for visual effect.ConclusionsIn the case study presented above, a bicycle stem was redesigned for AM with significant weight savings as a major result. The resulting weight reduction from 190 g to 140 g represents a 26% weight reduction. The presented case study exemplifies the intermediate role of design optimization tools that take advantage of AM design freedom, such as generative design in the present case. A key takeaway is that while GD might produce a design feasible to manufacture1993using an AM process, additional work, notably a rigorous and encompassing DFAM study, is required. Numerous practical challenges for the AM process and the subsequent manufacturing workflow persist which must be considered for a successful redesign.

Numerous practical challenges for the AM process and the subsequent manufacturing workflow persist which must be considered for a successful redesign. As with any redesign, a thorough economic analysis would be required to fully evaluate whether the weight savings realized through generative redesign are economically worthwhile. Although such an analysis is beyond the scope of the present work, the significant weight savings demonstrate the capacity of generative design to redesign light-weighted components.Specifically, the GD tool utilized in this work includes DFAM heuristics limited to build orientation, allowable overhang angle, and minimum feature size. However, as demonstrated in the examined case study, further considerations were necessary to fully prepare a design for the manufacturing process. Human intervention in the design process resulted in numerous other heuristics levied against the early embodiment designs, ultimately improving the product.

Human intervention in the design process resulted in numerous other heuristics levied against the early embodiment designs, ultimately improving the product. For example, DFAM heuristics were used to remove unbuildable geometry such as the large horizontal bores of the handlebar and steerer tube mounts and modify the geometry of the bolt hole bosses to avoid requiring supports. Other DFM inputs to the detail design phase were critical. Supports were designed to allow for efficient and cost-effective build plate separation. Rough, as-built surfaces were remedied after heat treatment but before machining. Machining allowances were added to critical surfaces, tool access and fixturing were considered, and order of set-ups designed.

Machining allowances were added to critical surfaces, tool access and fixturing were considered, and order of set-ups designed. These tasks represent restrictive DFM which goes far beyond simple DFAM heuristics and is based in a detailed consideration of the entire manufacturing workflow.In summary, the presented case study was used to demonstrate the multitude of practical challenges that arise when designing an AM part. Specific challenges related to generative design were overcome through human intervention and design modifications. Secondary processing steps were identified during the design process and modifications to the design were made to fluidly incorporate these steps. Through consideration of the restrictions and requirements of the necessary manufacturing processes throughout the entire manufacturing workflow, design modifications were made to improve the manufacturability of the proposed design.References[1] D.W.

chard F. Aubin United Technologies/Pratt & Whitney 400 Main Street MS 118-38 East Hartford, CT 06108 Phone 203-565-9073A CASE STUDY IN RAPID PROTOTYPINGINTRODUCTION:Since the first quarter of 1988, Pratt & Whitney (P&W), a Division of United Technologies Corporation (UTC), has been involved in the development for the use of rapid prototyping technologies that use a "layer-by-Iayer" building approach. Based on over 4 1/2 years experience with Stereolithography, this paper will address three aspects of our experience including: Implementation, Current Operations, and Future Plans.BACKGROUND:Component parts for gas turbine engines are required to function in environments that are unlike any other industry, with combustor temperatures reaching 2,500 degrees F., and pull loads on our 1st blades reaching 50,000 gls.

However, manufacturing these parts is accomplished using conventional methods that are similar to other industries.Conventional methods for prototype fabrication, small batch production and manufacturing process development have relied on drawings, hard tooling, numerical control programming, complex machining and/or extensive hand work. These approaches are typically characterized by long lead-time, high cost and multiple iterations to achieve desired results.There are many organizations today, that continue to employ these methods, but before they can make the transition into the new era of rapid prototyping, they must recognize that they are out of step with the current environment. Without feeling the need to change, the "Boiled Frog Phenomenon" is likely to occur.The Boiled Frog:This phenomenon is based on a classical experiment in biology.

Without feeling the need to change, the "Boiled Frog Phenomenon" is likely to occur.The Boiled Frog:This phenomenon is based on a classical experiment in biology. A frog which is placed in a pan of cold water but which still has the freedom to jump out can be boiled if the temperature change is gradual, for it is not aware of the barely detectable changing heat threshold. contrast, a frog dropped in a pot of boiling water will immediately jump out: it has felt the need to survive. changing environments, are likely to become "boiled frogs"; they act in ignorant bliss of environmental triggers and eventually are doomed to failure.In a similar vein, organizations and managers that are insensitive to gradually174InRichard United Technologies/Pratt & WhitneyAubinIMPLEMENTATION:Up until 1988, P&W was deeply involved in conventional prototyping methods.

We had an extensive investment in CAD/CAM equipment and trained personnel, but still had to "iterate" at the expense of extended lead times and increased costs. However, unlike the frog, we felt the water getting warm before it was too late.Along came an opportunity to "beta test" a new technology called Stereolithography (SLA) from 3D Systems Inc., of Valencia, California. Although faced with many unknowns and associated risks, we decided to get involved, and acquired 2 "beta machines" in January February 1988. One SLA machine with a 9-inch part building capacity was located at our facility in West Palm Beach, Florida, while the second one was located at our manufacturing facility in East Hartford, Connecticut.We were fortunate to have extensive resources in computer modeling hardware and software, and experienced people proficient in its use. computer models from the CAD systems to SLA, but learned quickly that complex computer models were very large in size.

computer models from the CAD systems to SLA, but learned quickly that complex computer models were very large in size. By using our shop computer network system (ethernet), we are now able to download computer models directly to the SLA machine.Initially, we used floppy disks to transport theTo increase the exposure of this new technology and to demystify its operation, we gave walking tours and presentations to:•Managers - So they understood its potential "fit" for their requirements, Engineers - So they understood the computer modeling requirements, e.g., bases &supports,..Mfg. Shop - So they could understand its capabilities and how to get access.We soon felt that we were a "tours-r-us" operation, but it paid off. Shortly, we were building a large variety of prototype objects and identified many tangible and intangible benefits.One of the most significant tangible benefits was the identification of design errors early in the design process.

Because SLA was so fast and relatively inexpensive to produce a prototype as compared to traditional methods, engineers were given more latitude to improve their designs, In the past, too many design so design quality was enhanced prior to reaching manufacturing. flaws were uncovered after the manufacturing departments made the tooling and the parts, investing time and resources only to find out they did the "wrong part right." By working out these errors before involving manufacturing, we can better align our shop resources to doing the "right part right."Joel Mokyr, Professor of Economics and History at Northwestern University and author ofhis recently published book titled, The Lever Of Riches, states: "Technological progress has been one ofthe most potent forces in history in that it has provided society with what175Richard F.

Aubin United Technologies/Pratt & Whitneyeconomists call a 'free lunch, ' that is, an increase in output that is not commensurate with the increase in effort and cost necessary to bring it about." There have been several occasions where, as a result of rapid prototyping, we have had a taste of this ''free lunch."Another tangible benefit was the improvement in communications among design engineers, and between engineering and manufacturing personnel because they now had a physical object to use instead of a sketch or drawing that was subject to interpretation.On the intangible benefits side, the SLA process provided an incentive and reward mechanism to our engineers to learn and use computer solid modeling that for more than just "pretty pictures".

By generating computer models of their design ideas, they would be rewarded with the satisfaction of possessing a physical prototype in return.Although considered intangible, our engineers were thus enabled to "think-out-loud", meaning that they were now able to generate physical objects directly from their computer models quickly and inexpensively. A new degree of enthusiasm resulted in the engineering community as a result of this new capability that essentially kept them "fresh" instead of growing "stale" while waiting for the hardware to arrive.Further, engineers were now receptive to change by incorporating suggested improvements from peers and from manufacturing personnel.Needless to say, these benefits provided significant time and cost savings.By the end of the first year of using SLA, we recognized the value of working together with other "beta test companies". Collectively, we worked and shared ideas to improve accuracy and surface finish.

Collectively, we worked and shared ideas to improve accuracy and surface finish. We also identified errors in the SLA process, and provided suggestions for improvement by submitting enhancement proposals to 3D Systems. Eventually, these enhancements were manifested in later revisions of the software, so mutual benefit was realized as a result of team-playing.With a few "free lunches" under our belt, we finished the first year with some valuable lessons learned, and unlike the "boiled frog," we found ourselves rewarded by being sensitive to our changing environment.After beta testing the early version of the SLA device, we were given another opportunity. This time, 3D Systems was offering their larger SLA machine for beta test, that had a 20-inch part building capacity.

This time, 3D Systems was offering their larger SLA machine for beta test, that had a 20-inch part building capacity. This gave P&W the capability to build larger prototype objects; we proceeded without hesitation.Having recognized the benefits of working with other companies throughout the beta testing programs, P&Wand the other beta test companies took the initiative to establish the SLA users group. This group was comprised of all users of SLA for the purpose of providing a forum for technical information exchange among the SLA users community. Through this forum,176Richard F. Aubin United Technologies/Pratt & Whitneycompanies share lessons learned, and identify strategies to improve their output.

Through this forum,176Richard F. Aubin United Technologies/Pratt & Whitneycompanies share lessons learned, and identify strategies to improve their output. P&W provided leadership to this group by chairing this organization from its inception through its first year.Internally, P&W decided to expand its rapid prototyping capabilities, and share its lessons learned with our sister divisions of UTC by establishing what we called an "Express Prototype Consortium." This consortium served as a vehicle for technology transfer with: Carrier, Hamilton Standard, Otis Elevator, Pratt & Whitney Canada, Sikorsky Aircraft, and United Technologies Automotive.CURRENT OPERATIONS:Currently, our operation in East Hartford has grown to include a total of 4 SLA machines and through the benefits of economies of scale, we are able to be responsive to the prototyping demands of our entire corporation.

Last year, 1990, over 1,600 prototypes were processed at our East Hartford facility.We see this activity in several basic areas:.. Design VerificationAs previously mentioned, there is great value in identifying design errors prior to manufacturing involvement. much more effective in time and cost than by conventional manufacturing techniques which iterate in "hard tooling."Improving designs by "iterating in plastic" is.. Manufacturing Producibility StudiesRapid prototyping is one of the key facilitators of integrated product development or concurrent engineering. Our manufacturing shops can get an earlier startup on process development by having a replica of a part to use to prove out tooling and fixturing while the actual real parts are being produced. For example, while investment castings are being processed, plastic replicas can be used in the manufacturing environment to ensure proper fixturing and tooling for final machining.

For example, while investment castings are being processed, plastic replicas can be used in the manufacturing environment to ensure proper fixturing and tooling for final machining. When the actual parts arrive, the machining process is already up and ready...Prototype Parts For Test And Conversion To Other MaterialsOccasionally, the plastic object can serve directly as a prototype in a test where the physical properties of the plastic can suffice. Techniques to convert the plastic object to other plastic materials having improved physical properties are also commonly available in industry through injection molding processes. And lastly, investment castings can be generated in a fraction of the time through the use of temporary tooling.177Richard F. Aubin United Technologies/Pratt & Whitneyit Research and Development EffortsTo date our biggest limitation with the photo-curable resins has been accuracy and surface finish.

Aubin United Technologies/Pratt & Whitneyit Research and Development EffortsTo date our biggest limitation with the photo-curable resins has been accuracy and surface finish. Our focus therefore has been to improve the process steps which affect them. We have learned that all variables in the total process must be considered, including: computer modeling preparation, hardware and software development, materials, and post-processing.Additional activities we have pursued include development relationships with resin suppliers to improve the basic resins themselves for less shrinkage, distortion, and better surface finish.Because of the benefits we realized from beta testing Stereolithography, we are currently involved in the beta testing of new emerging technologies:itSelective Laser Sintering (SLS), from DTM Corporation of Austin, Texas. This is a process that uses the heat of a C02 laser to sinter powdered materials in a layer-by-Iayer technique.

This is a process that uses the heat of a C02 laser to sinter powdered materials in a layer-by-Iayer technique. The materials offered include wax and poly carbonate plastic.itFused Deposition Modeling (FDM) from Stratasys Inc., of Minneapolis, Minnesota. This is a process that uses a heated extruder mechanism to deposit thin layers of wax or nylon like plastic materials.Both technologies build prototype objects directly from computer models, and each has its advantages; however, because of beta testing and non disclosure agreements, further information is prohibited at this time and is best derived directly from the companies developing the technologies.Finally, we are actively involved in several rapid prototyping consortia with academia including MIT, and the University of Texas.FUTURE PLANS:As you have seen, we have been very active players in the Stereolithography field.

We look forward to improvements in that field as it improves and in other technologies that we are hearing about in this conference. We feel strongly that continued development is essential to improve the accuracy and surface finish for prototype objects and to come closer to the "real" physical properties of "production-like parts. "As leaders in the application of rapid prototyping, we can't wait for these developments to mature. Although much can be done unilaterally and internally, more can be accomplished by pro-actively working with the manufacturers of rapid prototyping technologies, other users, and academia to drive development in the direction for maximum benefit to industry.The next "free lunch" just may evolve through dedicated continuous improvement.178.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCASTING - FORGING - MILLING COMPOSITE ADDITIVE MANUFACTURING THECHNOLOGYHaiou Zhang*, Runsheng Li*, Rui Wang*, Youheng Fu*, Xiangping Wang* , Guilan Wang†, and Shangyong Tang†State Key Laboratory of Digital Manufacturing Equipment and Technology, Huazhong University of Science and Technology,Wuhan, China, 430000 †State Key Laboratory of Materials Processing and Die and Mould Technology, Huazhong University of Science and Technology, Wuhan, China, 430000AbstractThe current metal additive manufacturing has some drawbacks, such as poor performance in producing forgings, low accuracy and efficiency, and high cost. This paper proposes a new technology called casting- forging-milling composite (CFMC) additive manufacturing, which uses the efficient and cheap arc as the heat source.

This paper proposes a new technology called casting- forging-milling composite (CFMC) additive manufacturing, which uses the efficient and cheap arc as the heat source. Synchronous arc welding and continuous semi-solid in-situ forging is achieved with micro-roller. Milling is incorporated into this process to remove defects and complete the part. Workpieces with equiaxed fine-grained microstructure and better performance than forgings have been obtained using CFMC. Testing shows that the mechanical performance exceeds the standards for forgings and most indicators are above the average levels. Trial products include titanium alloy aeronautical parts, a stainless steel propeller, and an aero- engine transition section which has passed the European standard x-ray inspection and test.

Trial products include titanium alloy aeronautical parts, a stainless steel propeller, and an aero- engine transition section which has passed the European standard x-ray inspection and test. The streamlined and low-cost manufacturing process achieved by using metal wire, integrated equipment and low pressure makes CFMC a green manufacturing model.IntroductionConventional machinery manufacturing has several serious weaknesses. First, it is highly dependent onoversize forging machines and dies, which require high initial investments and operating costs . Second, the conventional manufacturing processes are normally lengthy and characterized by separate casting, forging, welding and milling, repeated heating, huge energy consumption, and high pollutant emissions. Third, thin- walled parts can be manufactured only by subtractive manufacturing, which is a hugely wasteful method with a material utilization rate lower than 10%.

Third, thin- walled parts can be manufactured only by subtractive manufacturing, which is a hugely wasteful method with a material utilization rate lower than 10%. Fourth, it is difficult to manufacture functionally graded material (FGM) parts using conventional technologies [1].At present, metal additive manufacturing has a broad range of applications. There are three main additive manufacturing methods, which respectively use laser, electron beam, and arc as the heat source [2-5]. Additive manufacturing of high-performance parts not only seeks to produce forged parts with stable and reliable microstructures and properties, but also pursues high efficiency and low cost. However, the industrialization of large, high-performance forging production is bottlenecked by the following problems:High energy beam (e.g. laser beam, electron beam or arc) manufacturing is virtually casting by welding deposition and does not involve forging.

laser beam, electron beam or arc) manufacturing is virtually casting by welding deposition and does not involve forging. As high-performance parts are prone to metallurgical defects, warping and even cracking, large formed parts require high forming accuracy and can hardly attain desired mechanical properties for forgings when subject to dynamic loads. The anisotropy and unevenness of grain size, crystal orientation, and crystal structure cause great difficulty in internal structure and performance control [6-8].Unconstrained complex boundary and rapid alternate heating and cooling by a moving heat source cause uncertain and complex variations in thermodynamic conditions.

The uncertainty and low controllability of the growth, microstructure, morphology, and size distribution of grains can affect the performance of finished parts [9-10].The aforementioned three techniques of high-energy beam additive manufacturing have advantages in terms of forming accuracy, forming efficiency, and operating cost, respectively. However, none of them meets all requirements related to performance stability and the three factors. 1989The CFMC additive manufacturing technology proposed in this paper can ensure high forming accuracy, high efficiency and low costs while offering high performance [11]. During production, continuous rolling was applied to the arc deposition’s semi-solidification zone and finish milling was carried out one segment at a time [12]. Compared to laser additive manufacturing, CFMC additive manufacturing can offer higher deposition rate and formed parts with higher performance and stability at lower costs.

Compared to laser additive manufacturing, CFMC additive manufacturing can offer higher deposition rate and formed parts with higher performance and stability at lower costs. Compared with electron beam additive manufacturing, it not only has advantages in performance of formed parts and cost, but also removes the constraints on the dimensions of parts being manufactured imposed by the size of the vacuum chamber.Manufacturing processThe CFMC additive manufacturing process includes the following three steps: • Welding deposition • Micro-rolling • MillingStartSlice/Path planningCodeContinuous ForgingWelding DepositionAdjustment /MillingDefect?YesNoMillingEndFig. 1 Flow chart of CFMC additive manufacturingFirst, proper process parameters are determined by process analysis of parts. Fig. 1 shows the flowchart of CFMC additive manufacturing. Slicing and path planning for the model are conducted using computer software [13].

Fig. 1 shows the flowchart of CFMC additive manufacturing. Slicing and path planning for the model are conducted using computer software [13]. A movement mechanism with a welding torch, which is controlled by the forming path code, is employed to perform welding deposition. Appropriate values are assigned to the process parameters, including wire feed rate, voltage, current, etc. Then deposition by wire-fed arc welding is performed, with argon being used as a shielding gas [14]. During welding deposition, heat is rapidly transferred to the substrate and the microstructure of the molten metal then evolves into a dendritic crystal or columnar crystal structure [15].The second step is in-situ rolling, which uses a computer-controlled micro-rolling mechanism to apply a rolling force to the metal (Fig. 2). Then the metal undergoes plastic deformation and its columnar crystals break into fragments.

2). Then the metal undergoes plastic deformation and its columnar crystals break into fragments. Due to the effects of plastic deformation and heat input, the metal recrystallizes into equiaxed fine grains, and forgings are then obtained. Meanwhile, the on-line monitoring system performs data acquisition via the camera and data analysis and then feedbacks the results to the defect treatment module [16].Next, the defect treatment mechanism treats the defects by means of automatic adjustment of processparameters, milling, etc. Then milling is performed to improve forming accuracy and roughness [17].The composite manufacturing process employs two cooling sources: micro-roller and substrate. Normally, the first layer cools at the fastest rate and the rate of cooling decreases as the layer number increases.

Normally, the first layer cools at the fastest rate and the rate of cooling decreases as the layer number increases. In order to ensure the cooling rate does not vary across layers, additional cooling equipment is needed as a complement.1990This technology applies to a variety of metals, such as titanium alloys, stainless steels, nickel-based superalloys, austenite-bainite steels, and medium-carbon steel which has relatively low weldability [18].Fig. 2 Schematic of casting and forgingEquipment and devicesThe equipment used in this technology includes a gantry-type arc welding and casting-forging-milling system equipped with numerical control system, an automatic positioning system with two executable units and an external control shaft, an automatic tool changer, and a hydraulic micro-forging system.

The whole manufacturing system is highly flexible: it can perform welding deposition, forging, milling, and other operations, and the maximum dimensions and weight of workpiece it is capable of working are 5000mm*3000mm*2000mm and 15 tons, respectively (Fig. 3).To guarantee the stable performance and reliable quality of the high-end parts fab ricated by additive manufacturing, online nondestructive testing must be conducted during welding deposition to examine the welding parameters and surface morphology of welding bead. During forming, infrared thermography and ultrasonic testing can be used to identify the size, location, nature, and quantity of defects in workpiece, such as surface and near-surface cracks, lack of fusion, voids, and slag inclusions. The characteristics of arc column and shape of welding pool can be detected using a system for image capture and computer graphic analysis and diagnosis.

The characteristics of arc column and shape of welding pool can be detected using a system for image capture and computer graphic analysis and diagnosis. A real-time data acquisition system is used to monitor forming process parameters like arc current, voltage, track and velocity of motion of tool/workpiece, and wire feed rate [19].We have used cameras and infrared devices to track and monitor the whole process of CFMC additive manufacturing. Fig. 4 shows the thermograms produced. It is clear from the figure that the specimen’s temperature field had a uniform distribution [20].1991Fig. 3 Equipment for CFMC additive manufacturingFig. 4 Monitoring of CFMC additive manufacturing by infrared thermographyNumerical simulationThe stress variations in the freeform welding deposition (FWD) and Casting-Forging-Milling composite additive manufacturing technology (CFMC) specimens were simulated by ABAQUS. Fig. 5 shows the contours of maximum principal stress in the two specimens.

Fig. 5 shows the contours of maximum principal stress in the two specimens. It is clear from the figure that the maximum principal stress in the welding deposition specimens was primarily distributed in its central part and along the two long edges at the bottom. The CFMC additive manufacturing specimen had a greater surface area after forging and the deformation region created by forging mainly experienced compressive stress. The central part of the surface of the welding bead experienced compressive stress, while tensile stress occurred in the lateral parts. After forging, compressive stress arose in the region that was originally under tension after welding deposition. Vertically, plastic deformation occurred throughout the height of CFMC additive manufacturing.

Vertically, plastic deformation occurred throughout the height of CFMC additive manufacturing. Given that tensile stress is the major cause of crack initiation and growth, it is reasonable to infer that micro-rolling can effectively inhibit initiation of cracks.The evolution of the microstructures of the two specimens was modeled using MATLAB simulation program and the results are illustrated in Fig. 6 [21]. The rolling process increased dislocation density and induced dynamic recrystallization, which generated ultra-fine grains. This altered the developed columnar crystal structure resulting from welding deposition.1992Fig. 5 Contours of maximum principal stress in FWD and CFMC specimensFig. 6 Comparison of crystallization and crystal morphology in FWD and CFMC specimensExperimentsThe manufacturing process combining welding deposition with micro casting-forging was simulated using computer software. Then experiments were carried out to verify the simulation results.

Then experiments were carried out to verify the simulation results. In the experiments, the welding direction and deposition direction were set as the x- and z-axis directions, respectively, and the horizontal direction perpendicular to the x-axis was the y-axis direction (Fig. 7). Different processes were then selected for different types of metals. Metal-inert gas (MIG) welding was selected for medium-carbon steel, titanium alloy, and stainless steel, while tungsten inert gas (TIG) welding was applied to nickel-based superalloy. Table 1-3 shows the process parameters for different metal.Proper insert shielding gases were used to protect the welding areas. A protective cover or atmosphere room was needed to prevent oxidization of titanium alloys. Then suitable heat treatment processes were selected for titanium alloy and nickel-based superalloy.

A protective cover or atmosphere room was needed to prevent oxidization of titanium alloys. Then suitable heat treatment processes were selected for titanium alloy and nickel-based superalloy. Tensile testing at room temperature was performed on tensile specimens of medium-carbon steel, titanium alloy and nickel-based superalloy in accordance with the Chinese standard “Metallic Materials – Tensile testing – Part 1: Method of test at Room Temperature” (GB/T 228-2010), in order to measure their mechanical properties. Charpy impact tests at room temperature were conducted on specimens of titanium alloy 1993and nickel-based superalloy in accordance with the Chinese standard “Metallic materials – Charpy pendulum impact test method” (GB/T 229-2007).

Later, the medium-carbon steel specimens manufactured by freeform welding deposition (FWD specimen) and CFMC additive manufacturing (CFMC specimen) were scanned using a scanning electron microscope for their microstructures.monitorZWelding torchWelding beadWelding directionSubstrateXmicro-rollingYFig.

Figs. 8a and 8c show the morphology of crystals in the FWD specimen. The X-section image (Fig. 8a) reveals the initial distribution pattern of crystal grains along the height of deposition: crystals grew from the vicinity of fusion lines, along the direction indicated by maximum temperature gradient, towards to the central part, resulting in a columnar microstructure. Along the fusion line, the columnar crystals became coarser with increasing height. They were approximately perpendicular to the fusion lines. Temperature gradients were large during welding cooling. The grains that were oriented in the same direction as the maximum temperature gradient grew rapidly. As crystals oriented in other directions grew at slower rates, they met the adjacent rapidly-growing grains before they developed well, and thus their growth was restrained. Ultimately, the columnar crystals showed roughly the same orientation.

Ultimately, the columnar crystals showed roughly the same orientation. The coarse columnar crystals revealed in the figure were arranged in a dendritic pattern and had diameters of nearly 150 μm. As displayed by the Y-section of the FWD specimen (Fig. 8c), the initial crystal grains in different welding beads were largely columnar, but only those in the same welding bead were uniform in orientation, i.e. perpendicular to the corresponding fusion line. The orientation of columnar crystal in the Y- section varied between welding beads and the grains near each fusion line differed significantly in size.Figs. 8b and 8d show the morphology of grains in the CFMC specimen. In the X-section (Fig. 8b), the originally columnar crystals nearly disappeared, because most of them had been shaped by micro-rolling to equiaxed grains. Compared to the initial grains in the X-section of the FWD specimen (Fig. 8a), the crystal grains in the CFMC specimen had much smaller sizes.

Compared to the initial grains in the X-section of the FWD specimen (Fig. 8a), the crystal grains in the CFMC specimen had much smaller sizes. The image of the Y-section (Fig. 8d) reveals that the original crystal growth pattern in which the parallel grains were perpendicular to the fusion lines had been heavily damaged, leaving a few small-sized columnar grains locally in the CFMC specimen. While the crystal grains in the regions within about 200 μm of the fusion lines maintained their initial columnar shape, crystals outside these regions were all broken and some of them exhibited shear distortion. A comparison of the Y- section images of the two specimens also reveals finer grain sizes in the CFMC specimen.(a)FWD X-section (b) CFMC X-section(c)FWD Y-section (d) CFMC Y-sectionFig. 8 Comparison of crystal grains in the FWD and CFMC1995(a)FWD X-section (b) CFMC X-section(c)FWD Y-section (d) CFMC Y-sectionFig.

8 Comparison of crystal grains in the FWD and CFMC1995(a)FWD X-section (b) CFMC X-section(c)FWD Y-section (d) CFMC Y-sectionFig. 9 SEM images showing the microstructures of FWD and CFMC specimens (10000x)Furthermore, the X-section and Y-section of the specimens were scanned by SEM and interlayer spacing in the SEM images was measured with a micrometer, in order to analyze the distribution of carbide (Fig.9). In the X-section SEM image of the FWD specimen (Fig. 9a), the interlayer spacing largely fell within the range of 0.3 to 0.5 μm. In the Y-section SEM image of the FWD specimen (Fig. 9c), the interlayer spacing was mostly between 0.4 and 0.5 μm. The interlayer spacing criteria for distinguishing between pearlite, sorbite and troostite are as follows: pearlite: 150-450 nm; sorbite: 80-150 nm; and troostite: 30 to 80 nm. Therefore, the FWD specimen was composed primarily of pearlite, and the CFMC specimen was dominated by sorbate.

Therefore, the FWD specimen was composed primarily of pearlite, and the CFMC specimen was dominated by sorbate. No troostite was found.Titanium alloys: In arc welding, repeated heating can easily cause structural heterogeneity of welding beads. Therefore,proper heat treatment and deformation amount can play a significant role in ensuring structural homogeneity.Titanium has two crystal structures: hexagonal close-packed structure, or α form, and body-centered cubic structure, or β form. The former structure occurs at room temperature and then changes to the latter through at 882 ℃ or above as a result of phase change. The microstructures of titanium alloys split into the following types: α+β equiaxed microstructure, α+β duplex microstructure, and α+β basketweave microstructure, and Widmanstatten microstructure.The experimental research found that titanium alloy specimen that remained undeformed after annealing at 850℃ had Widmanstatten and locally basketweave microstructures.

The specimen that was deformed by rolling exhibited no α phase at crystal boundaries, but increased equiaxed α phase; its interior part showed a α+ basketweave + transformed β triplex microstructure, which can give the alloy good comprehensive performance.1996(a)FWD 50x (b) CFMC 50x(c)FWD 50x (d) CFMC 50xFig. 10 Comparison of crystalline phases of titanium alloysFig. 10 demonstrates that the undeformed specimen had a coarse columnar crystal microstructure and the columnar crystal had a thick α phase boundary. In terms of microstructure, the continuous α phase along this boundary contained parallel α clusters that grew inward, a typical Widmannstatten microstructure. The main contributor to this microstructure is the insufficient heat treatment of the β phase titanium alloy or inadequate amount of deformation.

The main contributor to this microstructure is the insufficient heat treatment of the β phase titanium alloy or inadequate amount of deformation. As each heating process during welding is equivalent to a post-heat treatment, the welding area’s temperature tended to exceed the point at which the β-phase transformation began, thus facilitating the formation of the Widmannstatten microstructure. In contrast, the temperature in the heat-affected region was far below the phase transformation point, allowing the crystal grains within it to continue growing. Therefore, the Widmannstatten microstructure was not distributed throughout undeformed specimen. The annealing at 850 ℃ only homogenized the interior microstructure, but did not remove the β-phase coarse grains that spread throughout the height of the specimen and the Widmannstatten microstructure.

Since the Widmannstatten microstructure has poor plasticity, it is necessary to remove it during heat treatment of titanium alloys.The CFMC specimen displayed finer equiaxed β-phase grains and no continuous α phase along grain boundary. Its microstructures included uniformly distributed equiaxed α and transformed β microstructures. The α microstructure outnumbered the transformed β microstructure and the latter showed smaller interlayer spacing. Overall, the CFMC specimen consisted primarily of duplex microstructures, which combine the advantages of equiaxed and basketweave microstructures and guarantee high plasticity and toughness.Nickel-based superalloys： Nickel-based superalloy wires were tested by arc welding deposition and CFMC separately. The specimens produced by the two processes were compared in terms of structural and mechanical properties. Fig.

The specimens produced by the two processes were compared in terms of structural and mechanical properties. Fig. 11 reveals that the dendritic crystal structure in the interior of the CFMC specimen was damaged and the slender columnar grains that were observed in the FWD specimen were replaced by equiaxed grains as a result of dynamic recrystallization.1997(a)FWD (b)CFMC Fig.

11 Comparison of crystalline phases of nickel-based superalloy specimens produced by FWD andCFMCMechanical propertiesmedium-carbon steelTable 4 Mechanical properties of medium-carbon steelSpecimenTensile strength σb (MPa)Yield strength σ0.2 (MPa)Elongation δ (%)Hardness (HBS)Shrinkage ψ (%)Impact toughness (J·mm-2)Standards for forgings 835 688 FWD 973 CFMC490382 6441013 18229-285290 301-30840464447.5The mechanical properties of forgings listed in the table above are sourced from the Chinese standard“Steel forgings for aviation” (HB 5024-89).The rolled parts outperformed those obtained by welding deposition in mechanical properties.

Compared to the standards for forgings, the rolled parts had significantly improved mechanical properties, with a 16.5% increase in, 31.4% increase in yield strength, 80% increase in elongation, 8.1% increase in hardness, 15% increase in shrinkage, and 8.0% increase in impact toughness.Mechanical properties of titanium alloy After annealing at 850 ℃, the titanium alloy attained optimal mechanical properties. As shown in the table below, the mechanical properties of the titanium alloy specimens created by FWD were above the standard values for forgings except yield strength. The specimens obtained from CFMC + heat treatment (annealing at 850 ℃) exhibited the best mechanical properties.

Their tensile strength, yield strength, elongation, and impact roughness increased 18.3%, 16.7%, 57%, and 61.1%, respectively, from those of forgings, and were up 8.5%, 17.8%, 37.7%, and 39.2% from the FWD specimens.Table 5 Mechanical properties of titanium alloySpecimenTensile strength σb (MPa)Yield strength σ0.2 (MPa)Elongation δ (%)Impact toughness (J·mm-2)Standards for forgings FWD CFMC + heat treatment895976 1059828820 9661011.4 15.73540.5 56.4The mechanical properties of forgings listed in the table above are sourced from the Chinese standard“Titanium and titanium alloy forgings” (GB/T 25137-2010).1998Table 6 Mechanical properties of nickel-based superalloySpecimenStandards for forgingsX-axis Y-axis Z-axisTensile strength σb (MPa) 1253 1241 1241Yield strength σ0.2 (MPa) 1034 1034 1034Elongation δ (%)12 10 6Shrinkage ψ (%)15 12 8Impact toughness (J·mm-2)39*FWDX-axis Y-axis Z-axis724 740 674446 450 44919 15 18- -43 41 42CFMC + heat treatmentX-axis Y-axis Z-axis1302 1343 13471108 1137 116516 17 1626 21 -51 56 53The mechanical properties of forgings presented above are sourced from the Chinese standard “GH4169 alloy bars, forgings and rings” (GB/T 25137-2010), except the value of impact roughness, which is provided by the standard “GH4169 alloy rods for aerospace”.The high tensile and yield strengths of the FWD specimens of nickel-based superalloy suggest relatively high plasticity.

Their average impact roughness was 7.7% higher than the standard value for forgings. The specimens obtained from CFMC + heat treatment showed substantial improvements in comprehensive mechanical properties. The heat treatment process included heating at 1090 ℃ for 1-2 hours, air cooling at +950 ℃ for 1 hour, air cooling at +72 ℃ for 8 hours, furnace cooling at a rate of 50 ℃/h to 620 ℃, and air cooling for 8 hours. In this process, the specimen went through homogenization, solution hardening, and two-stage aging treatment. After the heat treatment, the interior crystal grains in the CFMC specimens changed to aquiaxed grains of varying sizes through recrystallization. The recrystallization process was accompanied by crystal twinning.

The recrystallization process was accompanied by crystal twinning. Due to the presence of aquiaxed crystals, the specimens can attain excellent comprehensive mechanical performance with higher tensile strength, yield strength, toughness, and hardness than corresponding standards for forgings.1999Conclusions and future worksA transition piece for aviation made of medium-carbon steelFig. 12 A transition piece for aviation and its X-ray imagesFig. 12 shows a transition piece for aero-engines developed in cooperation with AECC Xi’an Aero-Engine Ltd. Due to their non-uniform wall thickness, transition pieces manufactured by the traditional investment casting are prone to cracking, loosening, and deformation and likely to have high porosities. As a result, the manufacturing process has an extremely low yield of less than 10%. The transition piece shown in the figure above is produced by the CFMC additive manufacturing technology proposed in this study.

The transition piece shown in the figure above is produced by the CFMC additive manufacturing technology proposed in this study. The X-ray testing reveals no flaw in it. This CFMC part is proven to be micro-grained and super-strong and its metallurgical quality meets Ukraine's strict standard for aero-engine testing.Titanium alloy aeronautical partFig. 13 Titanium alloy aeronautical part before milling2000Austenite-bainite steel frog for high-speed railways(a) austenite-bainite steel frog(b)Free annealing (c)Austempering and tempering (d)Rolling and heat treatmentFig. 14 An austenite-bainite steel frog and microstructures of steelThis frog is intended for switches on high-speed railways. It is manufactured from austenite-bainite steel by the CFMC additive manufacturing technology and has better properties than forgings. This type of frog has been applied in practice.

It is manufactured from austenite-bainite steel by the CFMC additive manufacturing technology and has better properties than forgings. This type of frog has been applied in practice. The figure above shows that the crystal grains resulting from rolling and heat treatment are finer than those from austempering and tempering.2001Other sample pieces(a) Stainless steel propeller(b) An aeronautical part from high-silicon aluminum alloy(c)A highly-skewed integrated impeller (d) Alloy drag plateFig.15 Sample piecesAdvantages: • It performs casting and forging simultaneously and can produce forgings with a hyperfine aquiaxed grainstructure.2002CFMC can reduce the anisotropy of parts’ interior structures and increase their toughness, which guaranteeuniformity of materials and thereby improve components’ fatigue life and reliability. It is able to produce not only thin-walled metal parts, but also parts with non-uniform wall thicknesses.

It is able to produce not only thin-walled metal parts, but also parts with non-uniform wall thicknesses. As the most accurate forging methods can only produce fine-grain metals, the conventional techniques of metal additive manufacturing are incapable of manufacturing forgings.Less or no use of huge traditional casting and forging equipment In traditional machinery manufacturing, casting, forging, welding, and milling units CFMC will combine the separate units for casting, forging, welding, and milling in conventional heavy- equipment manufacturing into an integrated manufacturing unit, thus achieving light-equipment manufacturing with shorter processes, higher energy efficiency, and lower pollutant production.

The computer-controlled forming paths help significantly reduce equipment investment and material cost.High rate of material utilization Using metal wires as the raw material, CFMC can guarantee a rate of material utilization higher than 70%.The prices of wire materials are about only 1/10 of those of powder materials, which are widely used in common additive manufacturing techniques that involve laser melting of powder.High efficiency and low cost of arc The most popular techniques of metal additive manufacturing all use laser and powder based methods.Lasers are the most expensive devices used in metal additive manufacturing and most lasers need to be imported.

For a given power level, using arc as the heat source can reduce cost ten-fold compared to the cost of laser-based manufacturing.Shorter production cycles Because CFMC can control parts’ dimensions, shape, microstructures and properties, it can offer a deposition rate of 5-15 kg/h and a double-wire deposition rate of 20-30 kg/h. Thus a record decrease of 60% in lead time can be achieved: while it takes more than three months to manufacture a two-ton metal forging in the past, CFMC needs only about 10 days.Energy-efficient green manufacturing As the production equipment needs a power supply of only 50 kW, its energy consumption per unit time isabout only 2‰ of a huge pressing machine’s.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCellular and Topology Optimization of Beams under Bending: An Experimental StudyArjun Gopal, Gaurav Parihar, McKay Holt, Tanner Stinson, Manasvi Sharma, Dhruv Bhate* Ira A. Fulton Schools of Engineering, Arizona State University, Tempe, AZ *Corresponding Author: dhruv.bhate@asu.edu (D. Bhate)AbstractDesign for Additive Manufacturing (AM) includes concepts such as cellular materials and topology optimization that combine the capabilities of advanced computational design with those of AM technologies that can realize them. There is however, limited experimental study of the relative benefits of these different approaches to design. This paper examines these two different approaches, specifically in the context of maximizing the flexural rigidity of a beam under bending, while minimizing its mass.

This paper examines these two different approaches, specifically in the context of maximizing the flexural rigidity of a beam under bending, while minimizing its mass. A total of 23 beams were designed using commercially available cellular design, and topology optimization software. The Selective Laser Sintering (SLS) process was used to manufacture these beams with Nylon 12, which were then tested per ASTM D790 three-point bend test standards. The effect of varying the size and shape of cells on the flexural rigidity was studied using 15 different cellular designs. These results were then compared to six different topology optimized beam designs, as well as three solid and hollow baseline beams.

These results were then compared to six different topology optimized beam designs, as well as three solid and hollow baseline beams. These preliminary findings suggest that topology optimized shapes underperform their cellular counterparts with regard to specific stiffness, and that stochastic cellular shapes deserve deeper study.IntroductionIn several applications, a main focus of design is towards volume and mass reduction while improving structural performance. Topology and cellular design optimization are two popular approaches used to achieve these goals. Topology optimization is a free form design approach that aims at finding optimal distributions of material within a domain for a given set of objectives and constraints [1]. It has therefore been known to lead to new and unanticipated designs that typically outperform conventional low weight designs [2, 3].

It has therefore been known to lead to new and unanticipated designs that typically outperform conventional low weight designs [2, 3]. Cellular structures consist of a unit cell (combination of material and space) that is repeated in space to create a larger structure. It is difficult or even impossible to manufacture these complex geometries using traditional manufacturing processes. Additive manufacturing helps us accomplish the manufacturing needs for topology and cellular optimized designs and the marriage of new design tools with AM process technologies has a large potential for impacting several industries.Studies have been done for manufacturing topology optimized structures using Additive Manufacturing processes. Cellular structures are another weight reduction solution which have been widely studied. The repeating lattice structures provide better structural properties in some contexts, when compared to baseline designs [4].

The repeating lattice structures provide better structural properties in some contexts, when compared to baseline designs [4]. The organic nature of the structures have extensive applications in the biomedical and aerospace industry for making implants and light- weight aircraft structures [5]. Using this technology, aerospace companies like Airbus and1877Bombardier have been able to reduce the weight of components significantly without compromising structural rigidity [5].A key question from a practical standpoint is as follows: how well do the predictions of design optimization software hold up in practice? This is the question this work seeks to address in a small way, by way of examining the performance of the software in the context of the 3-point bend test. Altair Inspire was used to perform topology optimization for maximizing the stiffness with varying mass percentages.

Altair Inspire was used to perform topology optimization for maximizing the stiffness with varying mass percentages. The objective of this study was to maximize the flexural rigidity of beams while minimizing relative density. The beams were printed using Selective Laser Sintering (SLS), which is an Additive Manufacturing (AM) processes used for polymers. The primary advantage of SLS is that it does not require support structures for printing which has multiple advantages including printing complex geometries with high resolution, reduced post processing time since there are no supports to be removed, high build rate and good mechanical properties [6]. Test were performed using a 3-point bend test setup. Topology optimization and cellular structure generation were used for designing 24 beams made out of Nylon 12. In the following section, the specific methods used to design these beams is discussed in more detail.

In the following section, the specific methods used to design these beams is discussed in more detail. This is followed by a discussion of the experimental setup for the 3-point bend test, and the results obtained from these tests. The paper concludes with a discussion of the conclusions and insights gained from this study.Design MethodsFor this study, three baseline beams were used for comparison in a 3 point bend setup: a solid cuboid, a hollow cuboid, and a hollow cylinder. For the hollow cylinder the inner radius is 0.67 times of outer radius, this is the optimum ratio for a cylindrical beam under bending [7]. For the hollow cuboid the thickness is 10% of the total width. The dimensions of the baseline beams are listed in Table 1.Table 1.

For the hollow cuboid the thickness is 10% of the total width. The dimensions of the baseline beams are listed in Table 1.Table 1. Baseline Design ParametersNameLength (in)Breadth (in)Height (in)Diameter (in)Solid Cuboid7.20.450.45N/AHollow Cylinder7.2N/AN/AOuter: 0.45 Inner: 0.30Hollow Cuboid7.20.450.45N/AIn addition to the baseline solid and hollow beams, two different commercial software packages were used for design optimization: SolidThinking’s Inspire software was used for topology optimization, and nTopology’s Element was used for cellular material design. User interfaces of both software packages are demonstrated in Figures 1a and 1b, respectively.

1Exportv[► R unCloseCance lAppl,(a)(b)Figure 1. (a) SolidThinking Inspire topology optimization user interface, and (b) Cellularstructure generation tab in Element nTopologyTopology OptimizationA total of six designs, shown in Table 2, were created with the Inspire software, beginning with a solid beam and a 150N applied load corresponding to the 3 point bend setup with roller supports. Exclusion zones were defined around contact of the load and the supports, so as to ensure regions available for contact were not subject to optimization. These beams varied in percentage of mass retained and whether the PolyNURBs tool was used to smoothen the geometry. This was done to study the effects of this smoothing on the resulting performance of the beam.1879....J0XNameTopOpt 1TopOpt 2PolyNurb 1PolyNurb 2PolyNurb 3Table 2.

This was done to study the effects of this smoothing on the resulting performance of the beam.1879....J0XNameTopOpt 1TopOpt 2PolyNurb 1PolyNurb 2PolyNurb 3Table 2. Topology Optimization Design ParametersMass Retained25302530351880PolyNURBNoNoYesYesYesPolyNurb 430YesCellular DesignCellular structures were created by importing the baseline solid cuboid into the nTopology Element software. A total of 14 cellular designs were created, with a 1mm thick sandwich panel added to the top and bottom – these are shown in Tables 3 and 4. The panels span the length of the beam including the overhang and have a width identical to the bounding box cross-sectional dimensions of each of the beams.

The panels span the length of the beam including the overhang and have a width identical to the bounding box cross-sectional dimensions of each of the beams. This panel was added due to the irregular surfaces of the topology and cellular structures to allow for a flat surface to make testing more accurate.Of the 14 cellular beams, eight had periodic tessellations with different unit cell shapes and member thicknesses as shown in Table 3. Two of these eight designs had periodic tessellation with tet octet edge unit cell and were optimized for our loading conditions using the Pro version of the Element software. The remaining six beams (from the original 14), were generated with stochastic lattices using various combinations of modifier points and surface modifiers which the designer to vary the size of the stochastic cells, and their distribution along the length and cross-section. The range of designs generated with this approach are shown in Table 4.

The range of designs generated with this approach are shown in Table 4. The seed value for the underlying stochastic algorithm that generates these shapes was kept constant for all stochastic designs (Figure 2).Table 3.

Stochastic structure generation tab in Element nTopology1883NameDense Bottom 1Dense Bottom 2Dense Bottom 3Dense CenterStochastic 1Stochastic 2Table 3. Stochastic Design ParametersMin DiameterMax Diameter32021021011013121884ModifierSurfacePointSurfacePointPointPointThe material used to fabricate the beams was Nylon 12 (PA2200) printed on an EOS Formiga P110. The tensile modulus, tensile strength and density were used as the input material properties for the beam optimization, and are summarized in Table 5.Table 5. Material properties of Nylon 12 used in this study [8]PropertyValueUnitTensile Modulus Z-Direction1650MPaTensile Strength Z-Direction42MPaFlexural Modulus1500MPaDensity930kg/m^3Experimental SetupLoad ing p i nFore ·SpecimenSupporting p insFigure 3.

3-point bend test setup per ASTM standards [9].In engineering mechanics, flexure or bending characterizes the behavior of a slender structural element subjected to an external load applied perpendicular to the longitudinal axis of the element. A flexure test produces tensile stress in the convex side of the specimen and compression stress in the concave side, this creates an area of shear stress along the midline. To ensure the primary failure comes from tensile or compression stress the shear stress must be minimized, this is done by controlling the span to depth ratio (span length divided by the height/width of the specimen). All beams were designed in this work for a span to depth ratio of 16:1 per ASTM D790-17 standards [10], and the test setup used is shown in Figure 4.

All beams were designed in this work for a span to depth ratio of 16:1 per ASTM D790-17 standards [10], and the test setup used is shown in Figure 4. Additionally,1885in concurrence with the test standard, the overhang length for the beam on either side of the supports was set to be 10% of the span length.Figure 4.Test setup used for this studyThe flexural test measures the force required to bend a beam under 3-point loading conditions. In this study, as the cross section of these beams are different from a uniform cross section for which an area moment of inertia I may be readily computed, an effective flexural rigidity (E*I*) is instead used as the primary criterion for comparison between beam designs. Thus, while there is no specific E (elastic modulus) and I value that is computable for these beams, the product E*I* may be assessed from relating beam deflection at an applied load to derive this measure.

The relative density is the density of the test beam by the density of a solid beam of identical dimensions and the relative flexural rigidity is the flexural rigidity of the test beam by that of a solid beam of identical dimensions.Results and AnalysisFigure 5 shows the load-displacement responses of all 23 beams – while comparing these curves is misleading, since each beam has a different mass, it is still noteworthy that the stochastic cellular beams show a wider range than the periodic ones, and that topology optimization solutions are confined to a relatively narrow range at the bottom. The solid box, as expected shows the highest load bearing capability, and is also among the stiffest and the hollow box and hollow cylinder show similar behavior.

The solid box, as expected shows the highest load bearing capability, and is also among the stiffest and the hollow box and hollow cylinder show similar behavior. Figures 6 through 8 are isolated load-displacement plots for each of the three design strategies employed in this work: topology optimization, periodic cellular materials, and stochastic cellular materials, respectively.1886120100Baseline Stach.as ,c Perlndlc (lattice) Topology Optimized80:;:- .D,. : ; 60Solid Box0 ..J40Hollow Cyflnd,r2000051 52.5335Displacement (inches)Figure 5. Load-displacement plots for all beams tested, color coded by high level design strategy2018PolyNurb 116PolyNurb 2... "' 0...I1412108PolyNurb 3TopOpt 1TopOpt 2642000.20.40.6 0 .8 Displacement (inches)1.21.41.6Figure 6.

Bending load-displacement response of stochastic cellular beams onlyTo enable meaningful comparisons between these beams, we must isolate the contribution of mass from the estimated quantities of interest – for this work, the two parameters of interest are flexural rigidity, and the maximum bending stress experienced in the beam. Figure 9 shows the effective flexural rigidity plotted as a function of relative density, where relative density is the mass of the designed beam divided by the mass of its solid bounding (enveloping) box. Therefore, the solid beam has a relative density of 1.800Color: Design Approach700Stochastic Cellular MaterialsSize600Interval500Baseline • Cellular - Period ic • Cellular - Stochastic • Topology OptimizedN° < E 400~ . wCl 0 C "0 <El* (Nm " 2)300200100~ - - All TopologyOptimized SolutionsG;J -e0.20.40.6 Relative Density0.81.0PageFigure 9.

wCl 0 C "0 <El* (Nm " 2)300200100~ - - All TopologyOptimized SolutionsG;J -e0.20.40.6 Relative Density0.81.0PageFigure 9. Effective flexural rigidity (E*I*) plotted as a function of relative density of the beam, with solid beam representing a relative density of 11888As can be seen in Figure 9, the topology optimized beams, despite being optimized with the objective of maximizing stiffness at different mass fractions, greatly underperform the beams populated with cellular structures. Further, the stochastic cellular materials once again display a wide range of results, including two designs that have rigidity measures far higher than the majority of periodic and stochastic designs. Plotting the maximum bending stress (estimated at the peak load experienced by the beam) against relative density, as shown in Figure 10, shows similar trends.

Plotting the maximum bending stress (estimated at the peak load experienced by the beam) against relative density, as shown in Figure 10, shows similar trends. The stochastic 2 design (see table 3), which is a graded stochastic geometry in a hollow box, showed an ability to withstand very high peak loads for its relative density in comparison to the other solutions.'===================== r ' ==i'L__ ______ _15000 ~Color: Design ApproachSizeInterval;;; s V, V,10000S tochastic 2Baseline • Cellular - Periodic • Cellular - Stochastic • Topology Optimized~ V, "' C 'ti C QI co X"' ~5000•• •Cl 0 C ,, <Max Bending Stress (psi)•All Topology Optimized Soluti ns0Map Sha~0.20.40.6 Relative Density0.81.0Freq IPageFigure 10.

Maximum bending stress (estimated at the peak load experienced by the beam) plotted as a function of relative density of the beamA final study of interest was to examine the effect of PolyNURBs on the bending response. PolyNURBs is a feature in the Inspire software that enables smoothing of faceted geometries that are typical of topology optimization software outputs. As can be seen in Table 2, the topology optimized designs selected for the study included 3 geometries that were smoothened with PolyNURBs, and 2 that were printed as optimized, with all facets intact. While this represents a very small sample size, a comparison of the geometric efficiency of the flexural rigidity was performed. Following Berger et al. [11], the geometric efficiency is estimated as a dimensionless ratio of the effective flexural rigidity discussed before, divided by the flexural rigidity of an equivalent solid beam, and expressed as (E*I*/EI).

As can be seen in Figure 11a, smoothing a faceted geometry does have the effect of increasing the geometric efficiency of its bending rigidity, and also its maximum bending stress, as seen in Figure 12a. However, when plotted with respect to relative density (Figures 11b and 12b), it is apparent that the improvements are essentially emerging from an addition of mass that is typical with smoothing applications. Thus we may1889conclude that smoothing functions do improve the solution (at least in the context of maximizing stiffness and peak load per unit mass) but do so only at the expense of increasing mass of the structure.U2 U0.160.15~0.140.12i 0.1 0V,Ji ;-- • w0.1~0.080.060.04PolyNurbDesignBTopOpt0.050.000.300.400.45r -0.SO Relative Density0.SSE*I /Esls • PolyNurb • TopOpt0.600.65(a)(b)Figure 11.

Geometric efficiency of flexural rigidity: (a) as a comparison between PolyNURBs and the faceted topology optimized solution, and (b) versus relative density250020002000;;; &V, V, ~ 1n 1500 Cl C: i, C:i 1000 ]"' "' X ~ 1000500PolyNurb • TopOptMax Bending Stress (psi)500 ...L _ ___ -::---:-1 -:-N:---:b- - - - - , - - --:;T=o"".pO~ pt: - - -~Poy urDesign0 0.300.350.450.50 Relative Density0.550.600.65Figure 12. Maximum bending stress (a) as a comparison between PolyNURBs and the faceted topology optimized solution, and (b) versus relative densityDiscussionTopology optimization and cellular materials represent two exciting avenues for design enabled by AM. However, much work is needed to validate the results of using these design18900.7(0.71software solutions, and identifying trade-offs and risks in implementing supposedly optimized designs.

However, much work is needed to validate the results of using these design18900.7(0.71software solutions, and identifying trade-offs and risks in implementing supposedly optimized designs. In this rather limited, and preliminary work, an effort has been made to demonstrate how the application of these software solutions to well-studied problems such as the 3-point bend test may elucidate practical aspects of implementing these optimization software without relying on more complex, component level design.The reader is cautioned against necessarily interpreting from this study that topology optimization solutions will underperform lattice based structures in bending – in fact there is computational evidence suggesting the opposite [12].

The use of two different software packages is also complicated by their own assumptions – however the experimental results do suggest that there is perhaps a loss of redundancy with topology optimization when these parts are fabricated and tested, that is not observed as much with cellular materials. Ultimately, this work hopes to make the case for greater experimental study of optimized geometries, be they via topoloAcknowledgementsThis work was conducted as part of a research project in the MFG 598: Design for Additive Manufacturing course offered in the Spring of 2019 at the Arizona State University.

The authors wish to acknowledge support from the Maricopa County Industrial Development Authority (MCIDA) and the Fulton Schools of Engineering at Arizona State University, where these parts were fabricated and tested, as well as Altair Engineering and nTopology for providing access to their software that was used to generate the designs in this paper.References[1] Liu, J., Gaynor, A. T., Chen, S., Kang, Z., Suresh, K., Takezawa, A. Albert. (2018, May 03). Current and future trends in topology optimization for additive manufacturing. Retrieved March 6, 2019, from https://link.springer.com/article/10.1007/s00158-018-1994-3 [2] Bendsøe, M. P., & Sigmund, O. (2013). Topology optimization: Theory, methods, and applications. Berlin: Springer. [3] Rashid, R., Masood, S. H., Duan, R., Palanisamy, S., Huang, X., & Rahman Rashid, R. A. (n.d.). Topology Optimisation of Additively Manufactured Lattice for Three-Point bend Test.

llular Structures for Optimal PerformanceSarah Engelbrecht1, Luis Folgar2, David W. Rosen1, Gary Schulberger2, Jim Williams21Georgia Institute of Technology Primary Contact: david.rosen@me.gatech.edu, 404-894-96682Paramount Industries, Inc. Primary Contact: Luis.Folgar@paramountind.com, 215.757.9611Abstractlattice structures, enable unprecedented stiffness and strength characteristics, for a given weight. New design and CAD technologies to construct cellular materials are presented in this paper. Such materials have very complex geometries, hence the need for additive manufacturing processes to produce them. A series of experiments was performed to build and test parts fabricated using Selective Laser Sintering. Variations in mechanical properties were quantified and related to processing conditions. Examples help illustrate the variety of applications of cellular materials in the aerospace, automotive, motorsports, energy, electronics, and related industries.

Examples help illustrate the variety of applications of cellular materials in the aerospace, automotive, motorsports, energy, electronics, and related industries. A software tool is being developed to enable users to design and construct parts with cellular structures.Cellular material structures, such as honeycombs andKeywords: cellular materials, conformal lattice structures, selective laser sintering, additive manufacturing.1INTRODUCTION1.1 Cellular MaterialsThe concept of designed cellular materials is motivated by the desire to put material only where it is needed for a specific application. From a mechanical engineering viewpoint, a key advantage offered by cellular materials is high strength accompanied by a relatively low mass. These materials can provide good energy absorption characteristics and good thermal and acoustic insulation properties as well [1]. Cellular materials include foams, honeycombs, lattices, and similar constructions.

Cellular materials include foams, honeycombs, lattices, and similar constructions. When the characteristic lengths of the cells are in the range of 0.1 to 10 mm, we refer to these materials as mesostructured materials. Mesostructured materials that are not produced using stochastic processes (e.g. foaming) are called designed cellular materials. In this paper, we focus on designed lattice materials.1Cellular materials have a biomimetic - or bio-inspired - origin. Many naturally occurring materials have a porous construction that enables them to be light, stiff or compliant, and multifunctional, for example to conduct heat well or to convey fluids and nutrients. Ashby and co-workers [7] provide a good motivation for the study of cellular materials: “When modern man builds large load-bearing structures, he uses dense solids; steel, concrete, glass. When nature does the same, she generally uses cellular materials; wood, bone, coral.

When nature does the same, she generally uses cellular materials; wood, bone, coral. There must be a reason for it.” In this context, the term “cellular” does not just mean that the material is composed of living cells; rather it refers to the construction using struts, webs, pores, and/or channels.In the past 10 years, the area of lattice materials has received considerable research attention due to their inherent advantages over foams in providing light, stiff, and strong materials [1]. Lattice structures tend to have geometry variations in three dimensions; some of our designs are shown in Figure 1. As pointed out in [5], the strength of foams scales as ρ1.5, whereas lattice structure strength scales as ρ, where ρ is the volumetric density of the material. As a result, lattices with a ρ = 0.1 are about 3 times stronger than a typical foam.

As a result, lattices with a ρ = 0.1 are about 3 times stronger than a typical foam. The strength differences lie in the nature of material deformation: the foam is governed by cell wall bending, while lattice elements stretch and compress. The examples in Fig. 1 utilize the octet-truss (shown on the left), but many other lattice structures have been developed and studied (e.g., kagome, Kelvin foam). We have developed methods for designing lattice mesostructure for parts [8,10] and have developed design-for-manufacturing rules for their fabrication in SL.1 1 10.5 0.5 0.50 0 0wd1 1 1d d db) skin with single layer of lattice structure b) skin with single layer of lattice structurew w w0.5 0.5 0.5uv0 0 0 0 0 00 0 0u u u0.5 0.5 0.51 1 1v v va) octet-truss a) octet-trussc) Skin with 2 layers of truss structure made in SL c) Skin with 2 layers of truss structure made in SLFigure 1.

Octet-truss unit cell and example parts with octet-truss mesostructures.1.2 Design for Additive ManufacturingDesign for manufacturing (DFM) has typically meant that designers should tailor their designs to eliminate manufacturing difficulties and minimize costs. However, the improvement of rapid prototyping, or Additive Manufacturing (AM), technologies provides an opportunity to re-think DFM to take advantage of the unique capabilities of these technologies. Several companies are now using AM technologies for production manufacturing. For example, Siemens, Phonak, Widex, and the other hearing aid manufacturers use selective laser sintering (SLS) and stereolithography (SL) machines to produce hearing aid shells, Align Technology uses stereolithography to fabricate molds for producing clear braces (“aligners”), and Boeing and its suppliers use SLS to produce ducts and similar parts for F-18 fighter jets.

In the first three cases,2AM machines enable one-off, custom manufacturing of 10’s to 100’s of thousands of parts. In the last case, AM technology enables low volume manufacturing and, at least as importantly, piece part reductions to greatly simplify product assembly. More generally, the unique capabilities of AM technologies enable new opportunities for customization, improvements in product performance, multi-functionality, and lower overall manufacturing costs.

More generally, the unique capabilities of AM technologies enable new opportunities for customization, improvements in product performance, multi-functionality, and lower overall manufacturing costs. These unique capabilities include:Shape complexity: it is possible to build virtually any shape, lot sizes of one are practical, customized geometries are achieved readily, and shape optimization is enabled.Material complexity: material can be processed one point, or one layer, at a time, enabling the manufacture of parts with complex material compositions and designed property gradients.Hierarchical complexity: hierarchical multi-scale structures can be designed and fabricated from the microstructure through geometric mesostructure (sizes in the millimeter range) to the part-scale macrostructure.In this paper, we cover two main topics.

First, we are interested in developing geometric construction methods that enable designers to take advantage of the shape complexity capabilities of AM processes. Specifically, we develop a method for constructing cellular materials that conform to the shapes of part surfaces; when restricted to lattice structures we call such constructs Conformal Lattice StructuresTM (CLS). Second, we quantify the variations in mechanical properties of CLSTM designs seen in parts fabricated by Selective Laser Sintering (SLS).2 PARAMOUNT AND GEORGIA TECH COLLABORATIONParamount and Georgia Tech first started collaborating in Fall 2006 to study conformal lattice structures and to explore their potential in aerospace and related industries.

Concisely stated, our objectives included 1) developing cellular structure design and manufacturing technologies, 2) identifying applications, 3) protecting intellectual property related to cellular structure design and manufacturing, and 4) commercializing these technologies.At the time, Georgia Tech had software and design technologies to generate conformal lattice structure that could be applied to simple surfaces. It was not possible to create CLS on arbitrarily complex surfaces or to fill arbitrary volumes with CLS. Paramount had extensive experience with SLS technologies and materials, but had not experimented with cellular structures to any large extent.In the first year, Paramount and GT studied the strength and stiffness characteristics of several types of lattice structures. We conducted many finite-element analyses of lattice structures under various loading conditions. We identified the Cantley truss [2] as a being very efficient for stiffness and strength.

We conducted many finite-element analyses of lattice structures under various loading conditions. We identified the Cantley truss [2] as a being very efficient for stiffness and strength. The octet was reasonably good for some loading conditions, particularly when bending was prevalent.In the second year, we embarked on a two-part approach. First, we wanted to develop software that could create CLS on any surface or to fill a complex-shaped volume. Second, we wanted to identify process parameters and conditions to enable the manufacture of CLS with SLS machines. In particular, it was important to identify how small lattice struts could be fabricated and to determine process conditions that enabled their fabrication. As a by-product of these efforts, we also have insight into the failure mechanisms of CLS that provide progressive,3gradual failure, instead of catastrophic fracture.

As a by-product of these efforts, we also have insight into the failure mechanisms of CLS that provide progressive,3gradual failure, instead of catastrophic fracture. Each of these two avenues of work will be reported in this paper.3 CLS DESIGN METHODThe basic idea of how cellular materials are created is presented here. Four example primitive cell types are shown in Figure 2, three of which are lattice structures and the fourth is a foam. These cell types are 2-dimensional for simplicity of presentation. The octet lattice in Figure 1a is an example 3-D cell type. Lattice structures consist of a set of struts (beams) that connect the nodes of the lattice.To generate the cellular designs in Figure 1, the primitive cell types must be mapped into a mesh. In 2-D, the mesh consists of a set of connected quadrilaterals. In 3-D the mesh consists of hexahedra (6-sided volume elements with planar sides).

In 2-D, the mesh consists of a set of connected quadrilaterals. In 3-D the mesh consists of hexahedra (6-sided volume elements with planar sides). The uniqueness of our work is our use of conformal cellular structures, rather than uniform “lattice block” materials, that can be used to stiffen or strengthen a complex, curved surface. To see the difference between conformal and uniform structures, Figure 3a is an example uniform lattice structure, while Figure 3b shows a conformal lattice. Meshes for uniform structures consist of cube elements in 3-D (squares in 2- D), while for conformal structures, the mesh elements are general hexahedra. We have developed a new algorithm for generating conformal meshes that are used to create conformal lattice and cellular structures. An older algorithm for generating conformal lattice structures based on a mapped meshing approach [10] has been updated significantly.

An older algorithm for generating conformal lattice structures based on a mapped meshing approach [10] has been updated significantly. For our CLS meshes, we prefer that mesh elements are as cubic as possible; i.e., are of uniform thickness and uniform size. Such meshes are typically not generated by the free meshing methods in finite-element analysis codes, while typical part geometries are too complex for mapped meshing methods.a)b)c)d)Figure 2. Cellular primitives: three lattice structures and one web structure.a) unifo rm lattice a) unifo rm latticeb) conform al lattice b) conform al latticeFigure 3. Uniform and conformal lattice structures.3.1 CLS Construction MethodThe overall method for generating conformal cellular structures is shown in Figure 4. It consists of two main steps, indicated by the shaded rectangles: computing 3D conformal mesh,4and populating the mesh with cells. Inputs and outputs of the steps are shown as ovals.

It consists of two main steps, indicated by the shaded rectangles: computing 3D conformal mesh,4and populating the mesh with cells. Inputs and outputs of the steps are shown as ovals. Each step is detailed below.The objective of the meshing algorithm is to generate a conformal hexahedral mesh into which cells from the cell library can be placed. One or more layers of cellular structure can be placed to support the part’s skin. The input to the algorithm may be a CAD solid model of the part, a surface model of the part, or a triangulated surface model of the part. The triangulated surface model is very common in the rapid prototyping industry; it is known as an STL model and consists of a collection of triangles that approximate the curved surfaces of the part CAD model.

A method of constructing solid and STL models of lattice structures was presented in [9]; the method utilized the conformal lattice generation algorithm from [10].18 1818 1816 1616 1614 1414 1412 12Z Z12 1210 1010 108 88 86 66 60 04 42 20 00 02 24 42 24 4Figure 4. Method for CLSTM construction.X X5 53.2 Construct 3D Conformal MeshThe algorithm to generate a 3-D conformal mesh is shown in Figure 5. The first step is to divide the part boundary into regions, each of which is relatively flat. It is easier to control the mesh generation method if regions are flat. We implement an absolute angular deviation measure between surface or triangle normal vectors to determine if that surface or triangle should be added to the region being generated. For the purposes of presentation, we will assume that a STL file was given. Then, one triangle is chosen as the first triangle of a region.

For the purposes of presentation, we will assume that a STL file was given. Then, one triangle is chosen as the first triangle of a region. The normal vector of each connected triangles is compared to the normal of the first triangle; if they differ by less than a given tolerance, the triangle is added to the region. As an example, the tolerance for the simple part in Figure 4 was chosen so that the model consists of two regions: the cylindrical surface and the planar surface.50 0For each region, three main steps are performed as indicated by lines 3, 4, 5 of Figure 5. The first of these steps is to compute the offset of the object boundary. An offset is a collection of points that are at a specified distance away from the starting surface (distance is called the offset distance). As an example, an offset of a circle is a circle that is concentric with the first circle.

As an example, an offset of a circle is a circle that is concentric with the first circle. If you want the offset of a circle of radius 10 inches centered at the origin, and the offset distance is 1 inch, you would get a circle of radius 11 inches also centered at the origin. The offset of the curve from Figure 3 is shown in Figure 6. The positive offset curve (shown with a long-dashed curve) is offset in the direction of the original curve’s normal vector. Generally speaking, a positive offset results in a larger object, while a negative offset results in a smaller object. Offsets are standard geometric modeling operations that are typically available in mechanical CAD systems. We use an offset method developed for tessellated part surfaces [3], but any offsetting method could be applied.Algorithm Construct Conformal Mesh Input: CAD or STL model, desired element size Output: hexahedral mesh that conforms to the outer surface of the given model 1.

Partition part model into relatively flat regions. 2. For each region, 3. compute offset of region boundary 4. construct tri-parameter volume 5. divide parameterized volume into hexahedra 6. End for each region. 7. Ensure region boundaries match.Figure 5. Algorithm for constructing a conformal mesh.The second step is to construct a tri-parameter volume between the original surface and its offset. Conceptually, this is simpler than it may sound. Typically, curves and surfaces in CAD systems are defined using parametric equations. For example, the curve in Figure 6 would be parameterized by one parameter, u, that varies from 0 to 1, as in the well known Bezier, B-spline, or NURBS curves. Surfaces are parameterized using two parameters. Volumes are parameterized using three parameters and, hence, are called tri-parametric solids. This step of the algorithm assumes that the original boundary surface of the object is parameterized.

Volumes are parameterized using three parameters and, hence, are called tri-parametric solids. This step of the algorithm assumes that the original boundary surface of the object is parameterized. The same parameterization is transferred to the offset surface. Finally, a lofting is constructed from the original surface to the offset surface. Parameterized equations and the operations of lofting and offsetting are all standard, well known geometric operations in the field.Positive Offset Positive OffsetHowever, a complication arises since the original boundary surface of the object consists of triangles. That the original boundary surface, which was assumed to be parameterized the paragraph in above actually is not parameterized. Hence, a parameterization must beis,Normal vector Normal vectorNegative Offset Negative OffsetOffset Offset distance distanceOriginal Surface Original SurfaceFigure 6.

Hence, a parameterization must beis,Normal vector Normal vectorNegative Offset Negative OffsetOffset Offset distance distanceOriginal Surface Original SurfaceFigure 6. Positive and negative offset curves of a curve.6developed, the details of which are presented in [6].The third step is to generate the conformal mesh by dividing the parameterized volume into individual elements (hexahedra). This is a very straightforward step. After selecting either the number of elements or their typical size, increments in each of the three parameters are computed, which are used to generate elements of the mesh. For example, if the increment in u is chosen to be 0.1, then 10 elements in the u direction will be generated, since (1 - 0) / 0.1 = 10. Note that the numerator of the expression (1 - 0) is the difference between the maximum and minimum u values.

Note that the numerator of the expression (1 - 0) is the difference between the maximum and minimum u values. By successively incrementing each of the three parameters that define the tri-parametric volume, mesh elements are created.The final step in Algorithm Construct Conformal Mesh (line 7) is to ensure that region boundaries match by adjusting node positions and by adding elements, if necessary. Since the regions are parametrized separately, the hexahedral elements may not match well. Nodes from neighboring regions may be moved and merged to achieve matched boundaries. Also, a series of hexahedral or tetrahedral elements may be added in gaps between meshes in neighboring regions [6].3.3 Construct CLSThe algorithm for the second step (populate mesh with cells) of the overall conformal cellular structure design method from Figure 4 is shown in Figure 7. One input to the algorithm is the conformal hexahedral mesh that was generated in the first step.

One input to the algorithm is the conformal hexahedral mesh that was generated in the first step. The other input is the cell types contained within a library. The first step is to partition the mesh elements into regions such that within each region the loading conditions are similar on each element. These need not be the same regions that were used for mesh construction. For each region, a cell type from the cell type library is selected to populate the mesh elements in that region. The idea is to match the region’s loading conditions to the cell type, such that the cell type is effective at supporting the loading conditions. In this manner, the resulting cellular structure is more likely to be lighter for a given level of stress or deflection. The dashed gray box surrounding the “Select Mesh Elements” and “Select Cell Types” operation boxes indicate that these operations may be performed concurrently or maybe performed sequentially, depending upon the designer’s preference.

Note that each of the Select operations may be performed automatically by an algorithm or may be performed by the designer directly.The final step in Figure 7, “Apply Selected Cell Types to Selected Mesh Elements,” is where the actual cellular geometric model is constructed. This operation has been called population of mesh elements earlier. This operation is straightforward. One simply maps a cell type into a mesh element. Since both the cell type and the mesh element are defined parametrically, a simple parametric mapping algorithm can be applied to directly construct cell geometry.To construct a STL or solid model of the CLS, additional geometric construction operations must be performed. We utilize the approach described in [9], where solid models of the half- struts incident at each node of the mesh are constructed using Boolean operations is a solid modeling system.

Then, each solid “node” is tessellated and the triangles are written to a STL file.7Conformal Conformal Conformal Conformal Hexahedral Mesh Hexahedral Mesh Hexahedral Mesh Hexahedral MeshCell Type Library Cell Type Library Cell Type Library Cell Type LibrarySelect Mesh Elements Select Mesh Elements Select Mesh Elements Select Mesh Elements for each Selected Cell Type for each Selected Cell Type for each Selected Cell Type for each Selected Cell TypeSelect Cell Types Select Cell Types Select Cell Types Select Cell TypesMesh Elements for Mesh Elements for Mesh Elements for Mesh Elements for each Cell Type each Cell Type each Cell Type each Cell TypeCell Types Cell Types Cell Types Cell TypesApply Selected Cell Types to Apply Selected Cell Types to Apply Selected Cell Types to Selected Mesh Elements Selected Mesh Elements Selected Mesh ElementsConformal Cellular Conformal Cellular Conformal Cellular Geometric Model Geometric Model Geometric ModelFigure 7.

Algorithm to compute the conformal cellular geometric model.The resulting conformal cellular geometric model can be subjected to optimization methods in order to reduce weight, increase strength, increase stiffness, achieve some other objective, or achieve some combination of objectives. Such methods have been applied by a number of research groups [4,11].4 EXAMPLE OF CLS CONSTRUCTION METHODSAs an example, a simple cylindrical surface will be stiffened by one layer of conformal lattice structure using the method presented in Section 3. Figure 8a shows the original cylindrical surface defined by a STL file. The surface is 2 units in diameter and 2.5 units long. Constructed lattice structure is intended to be 0.25 units on a side. With an angular deviation measure of 50 degrees, the surface is partitioned into three regions, as shown in Figure 8b. A quadrilateral mesh is created in each region, then the mesh boundaries are adjusted and merged.

A quadrilateral mesh is created in each region, then the mesh boundaries are adjusted and merged. The result is shown in Figure 8c, where the different colors indicate the different regions. After offsetting the surfaces and mesh, the tri-parameter volumes and hexahedral mesh can be constructed (Figure 8d). The final step is inserting lattice geometry into each mesh element, resulting in the CLS model shown in Figure 8e.5 SLS FABRICATIONThe concept of reinforcing components with CLS has an inherent fabrication challenge due to the complexity of the geometries and features that are impossible to manufacture via conventional fabrication process. Indirect casting techniques can generate some complex CLS from a master pattern, but the approach is very expensive, time consuming, and requires a lot of trials before an accurate part can be achieved.

identified a part on an unmanned air vehicle (UAV) with properties and a geometry suitable for reinforcement and optimization with CLS. The part has a characteristic saddle surface and serves as a hatch-cover on the UAV. The CAD files of this part were optimized using the design algorithms from Georgia Tech. Figure 9 illustrates the hatch- cover files ready for SLS fabrication after optimization with CLS.Figure 9. Paramount Industries Inc. Hatch-Cover Component Reinforced with CLSParamount Industries Inc. successfully fabricated the hatch-cover via SLS and developed internal processes to overcome the many significant challenges when considering SLS for the fabrication of CLS. The SLS process has 4 variables that are extremely critical when the requirement is repeatable mechanical performance from a truss-like structure made out of struts that are ~1-mm thick. The most critical variable is the spatial orientation of the CLS during fabrication.

The most critical variable is the spatial orientation of the CLS during fabrication. Any CLS has an optimal orientation in the x-y-z plane which yields the best mechanical response when compared to FEA models performed prior to fabrication. The next three most important variables of our SLS process are laser power, scan-speed or hatch speed, and scan-spacing or hatch distance. Other critical process variables include the material quality, heat distribution, laser beam profile, the control of scales-and-offsets, and outline settings.The SLS-CLS process was dialed-in by testing the effect of orientation on the mechanical properties of parts; this study used 64 samples. The orientation and position of each sample was recorded and the test analyses of the first 8 samples revealed the optimal orientation. Mechanical test results, however, showed significant inconsistencies which were determined to originate from multiple process variables.

Mechanical test results, however, showed significant inconsistencies which were determined to originate from multiple process variables. The parts were labeled according to their position (top or bottom, left or right, and front or back) so the first sample in Figure 10 “TLF” for example refers to the hatch-cover on the Top-Left-Front corner.The preliminary test results for the first 16 samples are presented in Figure 10. These test results show inconsistencies in the failure mode of the struts. The parts were fabricated in 2 layers with 4 parts per layer. The stress-strain curves for the compression loading samples show gradual drops during the loading cycle which relate to the fracture failure of a complete row of struts within the lattice structure. This failure mode indicates a variation of mechanical properties among different struts.

This failure mode indicates a variation of mechanical properties among different struts. Using infrared imaging equipment and thermocouple sensors it was determined that the build envelop and the machine had hot and cold spots that had to be improved before uniform properties could be achieved. It was also determined that the parts on top created thermal gradients that affected the mechanical properties of the parts at the bottom of the cake.10Figure 10. Hatch Cover Test Results Orientation A – No Process OptimizationFigure 11. Hatch Cover Test Results Orientation B – With Process OptimizationThe next sets of hatch-covers were fabricated after all the process parameters and variables were dialed-in, optimized, and measured. Figure 11 shows the mechanical results of the compression loading test after the process optimizations.

Figure 11 shows the mechanical results of the compression loading test after the process optimizations. The significant difference is not only the shape of the stress-strain curve during the compression loading cycle and the repeatability of results, but also the maximum flexural load before failure going from 120 psi on results from Figure 10 to about 270 psi after optimization of orientation and processing parameters.116 CONCLUSIONSMethods to construct cellular material structures, specifically conformal lattice structures, were presented in this paper. Their fabrication using SLS was described and it was shown that process conditions and settings have a significant impact on mechanical properties of fabricated structures.

Based on this work, the following conclusions can be made:the presented CLS construction methods produced lattice structures on a wide variety of surface shapes, demonstrating the generality of the methods;these methods produced CLS that was of high quality, as measured by aspect ratios, skew, and related metrics, as well as of nearly uniform size;the geometric construction methods resulted in standard STL files that have been built in several AM machines, including stereolithography and SLS machines;mechanical properties of CLS are highly dependent on appropriate part orientation, as well as process variables of laser power, scan speed, scan spacing, and several others; by fine-tuning these variables it is possible to repeatedly build arrays of CLS parts with consistent, good mechanical properties.Future work includes converting the CLS construction software from Matlab to C++ and integrating the software with commercial CAD systems so that it can be disseminated and commercialized easily.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCERAMIC ADDITIVE MANUFACTURING: A REVIEW OF CURRENT STATUS AND CHALLENGESLi Yang, *Hadi MiyanajiDepartment of Industrial Engineering, University of Louisville, Louisville, KY 40292AbstractIn recent years, various additive manufacturing (AM) technologies that are capable of processing ceramic materials have been demonstrated. On one hand, many of the AM ceramic technologies have demonstrated geometry freedom capability and broad range of material flexibility. In some of the ceramic AM processes the part accuracies have also been favorably demonstrated. On the other hand, when reviewing the requirements of ceramic structures from applications perspective, there still appears to exist a misalignment between the demonstrated capability of ceramic AM and the required performance.

The lack of critical microstructural characteristic and performance evaluation results are likely setting significant barriers for the broader adoption of ceramic AM, which should be addressed via close collaborations between academia and industries.Keywords: Ceramics, additive manufacturing, review, applicationsIntroductionCompared to both polymers and metals, ceramic materials possess many attractive physical and mechanical properties, most notably the generally superior mechanical properties (e.g. modulus, hardness, strength) at elevated temperatures and many unique physical properties such as electrical and thermal conductivities that can be tailored at broad ranges for different applications [1]. In many application areas such as aerospace, automobile and energy, many structures and components are often subjected to high operation temperatures, such as the gas turbines, engines, batteries and heat exchangers, for which ceramic materials appear to be most suitable [2-5].

However, the structures for these applications are often complex due to various performance and assembly needs, which presents significant challenges to the ceramic materials. It is well known that ceramic materials generally possess low toughness, low ductility and high crack damage sensitivity, which, combined with the high hardness and high melting temperatures, make them difficult to work with during the manufacturing [6-8]. Furthermore, due to the intrinsic material characteristics of ceramics, multi-material systems are often needed in order to develop functional components, which further adds challenges to manufacturing [9-13].A vast body of literature exists about the manufacturing of ceramic materials, and it is rather difficult to establish a rigorous classification system to clearly distinguish all these manufacturing processes. In one of the recent works, it was proposed that these technologies can be classified into five categories [14]:1.

In one of the recent works, it was proposed that these technologies can be classified into five categories [14]:1. Casting/solidification process: Involves liquid-solid state change of the starting materials, which is often accompanied by volumetric change;2. Deformation process: Involves the shaping of ceramic structures via plastic deformation6523. Machining/Material removal: Involves the material removal of a ceramic workpiece using subtractive methods;4. Joining: Assembly of ceramic components via the creation of interfacial bonding; 5. Solid freeforming: involves additively manufacture the ceramics. Most of the traditional ceramic processing technologies fall into the first four categories. The casting/solidification processing encompasses a large number of processes including the suspension and gel based casting processes, the injection molding processes, and reaction based processes.

The casting/solidification processing encompasses a large number of processes including the suspension and gel based casting processes, the injection molding processes, and reaction based processes. Due to the high melting temperatures of ceramic materials, the casting and molding of ceramics are usually achieved via the phase transformation of the “carrier” phases, which can be either solvent solutions or matrix materials. In addition, most of the casting/solidification processes utilize defined surfaces (e.g. mold cavities, containers) to facilitate the shaping of the ceramic geometries. Therefore, in general these processes face multiple challenges. First of all, due to the state change of the carrier phases during the fabrication processes, the ceramic structures are often prone to defects such as voids and cracks [15-17]. Secondly, the loss of the carrier phases (e.g. evaporation) causes the potential loss of the geometrical accuracy [18, 19].

Secondly, the loss of the carrier phases (e.g. evaporation) causes the potential loss of the geometrical accuracy [18, 19]. Thirdly, the use of molds introduce significant limitations with geometrical design of the structures and render certain designs such as internal features infeasible [20, 21]. The deformation processing technologies include processes such as extrusion, gel foaming and superplastic forming, which mostly deal with ceramic suspensions and slurries that have higher viscosities, which enables them to be processed in similar ways as typical polymer materials [22, 23]. On the other hand, due to the extensive effects of shear, macro-phase segregation might occur, which in certain cases could even cause catastrophic part failure [24]. In addition, for feedstocks with non-spherical ceramic phases, the shear flow could also introduce additional anisotropy in the final material properties by aligning the ceramic phases preferably along the shear flow direction [24].

Compared to the first two categories, machining/material removal processing is relatively less broadly employed. This is largely due to the low damage tolerance and high crack sensitivity of the ceramic materials upon the machining. In many traditional ceramic material removal processes such as grinding and machining, damages such as pulverization and micro-cracking occur [25]. On the other hand, non- contact machining methods such as ultrasonic machining possess capability to process hard ceramic materials with minimum surface damage at the cost of processing speed, while laser machining suffers from the thermal residual stress due to the induced heat during the machining [25, 26]. A common approach to alleviate the machining damage of ceramics is to process the ceramic parts in their green states.

A common approach to alleviate the machining damage of ceramics is to process the ceramic parts in their green states. However, even for the machining of the green ceramic parts, micro-cracking is still likely to occur due to the loss of mechanical strength of exposed green part surfaces, which could only be partially alleviated even with the introduction of additional binder phases or pre-sintering processes [25, 27]. Due to the limitations with geometrical flexibility for processes of the first three categories, the joining processes such as ultrasonic welding, co-sintering and diffusion bonding are often employed. The use of joining for complex ceramic structures such as porous scaffolds [28], multi-material composites [29] and functionally graded materials [30, 31] have been demonstrated, which often involve rather extensive and labor-intensive process chains.

653including material extrusion, stereolithography, binder jetting, ink jet printing, powder bed fusion and sheet lamination. However, unlike the AM polymer and metal materials that have been widely used in various applications, there have been very limited use of AM ceramics due to various reasons. In this paper the authors attempt to perform a review with the current status of the AM ceramic fabrication in the attempt to reveal some of the critical developmental gaps in this area.Current applications of ceramicsWhile there exist various applications for ceramics, not all of them are readily justifiable for AM adoption. Many of these applications exploit only single attributes of the ceramic materials and do not require complex material/structural design. Such examples include the refractory containers, grinding wheels and chemical containers. With these applications, the benefit of AM is currently marginal.

Such examples include the refractory containers, grinding wheels and chemical containers. With these applications, the benefit of AM is currently marginal. On the other hand, for the areas in which multi-objective designs or integrated designs could bring about improved performance, such argument is more easily made. These structures often favor multi-material designs or structures with complex architectures that renders traditional manufacturing approaches problematic. With this in mind, three areas of applications were reviewed as promising future directions, which are the engine and propulsion, the dentistry and the electronics.Engine and propulsionOne of the most commonly exploited attributes of ceramic materials is their high- temperature mechanical properties, which allow them to survive the operations in extreme environments more reliably. Engine and propulsion components for aerospace, automobile and energy are among the most extensively investigation applications [32-35].

Engine and propulsion components for aerospace, automobile and energy are among the most extensively investigation applications [32-35]. In these applications, the push for ever-higher operation temperatures in the attempt to improve the performance efficiency has stretched the traditional superalloys to their limit [36]. For example, for the next generation gas turbine engines with operation temperatures of over 1200°C or so, ceramics are considered to be the only suitable material options [35]. In these applications, ceramics are primarily used in two ways: either as thermal barrier coating (TBC) or as primary structural materials. [35, 37, 38].The typical microstructural architectures of the TBC is shown in Fig.1, which clearly shows the TBC, the underlying superalloy substrate and the intermediate thermally grown oxide (TGO) layer in between [38].

The TBC topcoats usually have thickness of 100-2000µm and possess very low thermal conductivity across the entire operational temperature ranges in order to avoid the softening or melting of the superalloy substrates [38]. In addition, due to the frequent thermal cycles with large temperature ranges (room temperature to ~1300°C), the TBCs are also subjected to cyclic thermally-induced stress without fracture or delamination from the substrate, which usually have different coefficients of thermal expansion [35]. Such demanding requirements with both thermal and mechanical properties makes it difficult to find the suitable material candidates. Currently the most commonly utilized ceramic material for the TBC is the yttria stabilized zirconia 4+O7) and fluorite (YSZ), although some other potential candidates such as pyrochlores (A2 oxides have also been investigated [35, 37-40].

The 7%wt yttria stabilized zirconia (7YSZ) is made into porous structures with microstructural defects to introduce low thermal conductivity and high compliance [36]. The ferroelastic toughening mechanism unique to this composition of YSZ also makes it mechanically robust [37]. However, YSZ-based TBCs are not without drawbacks. For example, at elevated temperature the oxygen diffusion from YSZ to the underlying superalloy3+B2654substrates could significantly impact the performance of the structure, which necessitates the introduction of the TGO layer. The TGO layer usually consists of Al2O3 or mullite, which has very low oxygen diffusivity and bonds well with both the Ni-base superalloy and the TBCs [37, 38].

The TGO layer usually consists of Al2O3 or mullite, which has very low oxygen diffusivity and bonds well with both the Ni-base superalloy and the TBCs [37, 38]. In order to overcome some of the other limitations encountered by YSZ-based TBCs such as the limited melting temperature (~1300°C) [36], the degradation of compliance due to the infiltration of the calcium-magnesium alumino-silicate (CMAS) that is formed from the ingested environmental dust [36, 41, 42], and the insufficiently low thermal conductivity at even higher service temperatures [36], multi-layer surface coatings have been investigated, which aims to further enhance the performance of the TBCs via combining the advantages of properties from multiple layers [40, 41].

For example, due to its higher temperature stability and low susceptibility to the CMAS infiltration, the Gd2Zr2O7 (GZ) can be used as the topcoat, which is backed by the traditional YSZ that provides both toughness and good bonding with the TGO layer as shown in Fig.2 [40].Fig.1 Cross section of a YSZ TBC deposited on a superalloy [38]Fig.2 Cross section of a multi-layer GdZrO2/YSZ TBC on Hastelloy-X substrate [40]Currently the TBCs are usually applied to the substrate via various surface deposition methods such as electron beam physical vapor deposition [39, 40], suspension plasma spray [41, 43], suspension high velocity oxy-fuel spray [43] and air plasma spray [44]. Some of the common objectives for all the processes include the introduction of controlled porosities and porosity distributions [44, 45], and the control of microstructure of the TBC layers [46, 47], which is often achieved via both process parameter adjustment and substrate preparation.

Due to the layerwise structure and the need for controlled porosity for the fabrication of the TBCs, it appears that there exist potentials for the utilization of AM technologies. For example, it is known that the powder bed based AM processes possess the potential to produce samples with varying levels of porosity, which is often considered as defects but could be intentionally taken advantage of for the TBC fabrication.

On the other hand, since the coatings must be applied to existing surfaces, the only AM technology currently capable of the coating of non-flat substrate is the directed energy deposition process, which has some significant drawback for this particular application, as it generally lacks porosity control and also introduces large in-process thermal gradients that could cause horizontal cracks that are detrimental to the performance [43, 48].When even more stringent temperature requirement is imposed, the superalloy material could no longer meet the performance requirement, and the primary structures must also be made of ceramic materials.

In most of these applications, monolithic oxide and carbide ceramics are usually not used due to their intrinsic crack sensitivity and brittle fracture modes [33], although some of these materials such as HfB2 and ZrB2 have been investigated for ultra-high temperature applications (>1500°C) such as the leading edge for hypersonic aircraft, in which the extremely- high temperature oxidation resistance and thermal cycling lifetime become of critical importance655[34]. Currently the ceramic matrix composites (CMC) are most employed for these applications due to their potentials to combine the thermal properties of ceramics with the mechanical properties of the reinforcement phases, which are usually in fiber form in order to provide the maximum reinforcement effects [49].

This is necessary since beside the temperature requirements, these components are also subjected to various mechanical requirements such as creep resistance [33], damage tolerance [36], high temperature strength retaining [32], impact resistance [50] and lightweighting [51]. Due to the exposure to the liquid propellants and other contaminants, chemical resistance is also a critical requirement [52]. Typical CMCs used in engine propulsion comprise SiC fibers embedded in SiC matrix, although other combinations such as ZrB2/SiC [4], C/SiC [52] and Al2O3/Al2O3-LaPO4 [53] have also been reported. Fig.3 shows the comparison of mechanical and thermal capabilities of various ceramic matrix composites.

Fig.3 shows the comparison of mechanical and thermal capabilities of various ceramic matrix composites. As a matrix material, SiC possesses several attractive characteristics for such applications, including high melting point (~2500°C), good mechanical properties (hardness, strength, creep resistance) at high temperatures, relatively good oxidation resistance, and relatively high thermal conductivity [51, 54]. However, since the oxidation resistance of SiC originates from the formation of the surface oxide SiO2 layer, the material is prone to oxygen diffusion-induced oxidation failure, which could take place either through the matrix-reinforcement interphases or via the cracks upon damage [54, 55]. This issue imposes an important limiting factor during the design of both the reinforcement and interphase materials and the manufacturing processes.

This issue imposes an important limiting factor during the design of both the reinforcement and interphase materials and the manufacturing processes. For example, although carbon fiber could theoretically offer maximum high-temperature mechanical strength benefits, due to its intrinsic tendency of oxidation, it would introduce significant oxygen diffusion pathway to the SiC matrix [36, 54]. Similarly, in the design of the interphase, the bonding between the interphase and the fiber surface should be strong in order to avoid oxygen diffusion through the reinforcement fiber [5.9]. On the other hand, the interphases are often designed to intentionally introduce controlled crack propagation mechanisms in order to enhance the toughness and damage tolerance of the ceramic matrix, which is usually relatively limited [54].

As shown in Fig.4, the interlayers are often designed to have laminated micro-architectures or highly porous microstructure in order to direct the micro-crack propagation in the direction that is parallel to the fiber orientation.a. Temperature capabilityb. Mechanical strengthFig.3 Comparison of various ceramic matrix composites (and Ni-based superalloys) [36]a. Single-layer pyrocarbon or hexagonal BN Fig.4 Microstructural designs of interphases for ceramic matrix composites [54]c. Multilayer X-Y (X: pyrocarbon, Y: hexagonal BN)b. Porous SiC single layer656Multiple manufacturing processes are commonly employed to manufacture ceramic matrix composites, including liquid polymer infiltration [32], chemical vapor infiltration [33, 51], polymer impregnation and pyrolysis [33, 51], reactive melt infiltration [51], and impregnation- sintering [51].

In general the process routes for the ceramic matrix composites are similar to those for the other composites and are very complex, which makes the control of process quality a challenging task. This might have contributed to the very limited application of ceramic matrix composites in propulsion despite decades of intensive research and development. However, based on both the material and the structural requirements, it does not appear that AM currently has the capability to handle these materials, as there is a complete lack of literature with this subject. Multiple AM systems that are capable of fabricating fiber-reinforced composite structures have been developed, but they still lack the capability to deal with ceramic matrix materials [56, 57]. Due to the layerwise fabrication characteristics or AM, fiber reinforcements could only be oriented in the X-Y plane that is normal to the build direction.

Due to the layerwise fabrication characteristics or AM, fiber reinforcements could only be oriented in the X-Y plane that is normal to the build direction. On the other hand, in theory a pick-and- place mechanism can be used to place fibers into the newly formed surface, which can be (e.g. melting, the matrix via either subsequently embedded photopolymerization) or post-process treatment (e.g. pyrolysis, sintering, infiltration) [58, 59]. To partially offset the challenge of processing ceramic materials directly, it has been suggested that the potential of polymer precursors could potentially be taken further advantage of [58], which was demonstrated in a recent work of AM ceramic matrix composite by 3Dynamics [60]. In this work, siloxane-based polymers (e.g.

In this work, siloxane-based polymers (e.g. (mercaptopropyl)methylsiloxane and vinylmethoxysiloxane) and silazanes, which in combination with UV free-radical photoinitiator, free-radical inhibitor and UV absorber, were used as the photopolymer resin in the stereolithography-based process to fabricate precursor green parts. Further adding ceramic microfibers allows for the fabrication of CMC precursor, which after pyrolysis at 1000°C-1300°C in argon transforms to the CMC component based on silicon oxycarbonitride (SiOCN) matrix and ceramic microfiber reinforcement. Additives such as boron, zirconium were considered for the enhancement of high temperature capabilities of the structure, and additive such as silane could reduce the oxygen in the ceramic and pushes the ceramic composition towards SiC. However, for AM ceramic matrix composites, porosity and surface crack control remain significant challenges.

However, for AM ceramic matrix composites, porosity and surface crack control remain significant challenges. Due to the anisotropy of the matrix structures as a result of the layerwise process, uniform shrinkage is yet to be achieved for pyrolysis. In addition, this work only demonstrated the manufacturing of ceramic matrix composites with microfiber reinforcements, while long-fiber reinforcements remain out of limit.intoin-situ bondingDentistryCeramics have been widely used in dental applications for restorations of veneers, crowns and bridges [61, 62]. The most commonly investigated ceramic materials include alumina, leucite porcelain, zirconia and lithium disilicate. These materials are chosen largely due to their natural- looking characteristics, which are determined by various optical attributes such as color, opacity and translucency [63].

These materials are chosen largely due to their natural- looking characteristics, which are determined by various optical attributes such as color, opacity and translucency [63]. The dental ceramic materials are either used in combination with metals or in all-ceramic restorations, as illustrated in Fig.5 using dental crown restoration as an example. In veneer restorations, porcelain glass is among the first ceramics introduced into such applications, although it was not widely used before the concept of alumina (Al2O3) addition was introduced later in 1960s [62]. Another type of glass ceramic is the leucite reinforced glass ceramics. Leucite (KAlSi2O6) is a potassium alumino-silicate that has a tetragonal crystalline structure at room temperature and a cubic crystalline structure at temperature above 625°C. As a result, during the phase transformation, Leucite undergoes large volume change.

As a result, during the phase transformation, Leucite undergoes large volume change. Therefore, in the Leucite reinforced657glass ceramic systems, the SiO2-Al2O3-K2O system develops internal compressive stress upon cooling from the sintering/casting temperature, resulted in a strengthening effect that acts to inhibit the crack propagation within the ceramic structures [61, 64]. In comparison, for the all-ceramic systems, due to the intrinsic crack sensitivity of ceramic materials, the overall performance is still yet to match that of the metal-ceramic systems [65]. Among the all-ceramic materials, zirconia is currently most broadly used due to its combination of high toughness, low thermal conductivity, low corrosion potential, high biocompatibility and good radiopacity [62, 66, 67]. Yttria stabilized zirconia (YSZ) is most often used in order to improve the material processibility and to enhance the toughness.

Yttria stabilized zirconia (YSZ) is most often used in order to improve the material processibility and to enhance the toughness. Despite the promising properties, in clinical practice it has been found that dental parts made by zirconia or a combination of zirconia and layering glass ceramic have a higher incidence of fracture compared to prostheses fabricated exclusively out of dental metal alloys. One major cause of the failure is the occlusal overload due to bruxism, which causes crack at the cementation surface that propagates subsequently radically towards the surface and eventually causes the fracture of the entire restorations [65]. Also, there are other common minor complications that are related to the repeated loading of the restorations. Another issue with zirconia and other oxide ceramic materials is the low translucency and therefore undesired aesthetic properties.

Another issue with zirconia and other oxide ceramic materials is the low translucency and therefore undesired aesthetic properties. In order to overcome this, a top layer made from porcelain glass is often applied to the oxide ceramic restorations, adding another level of complexity to both process and quality assurance.a. Ceramic-on-metal b. All-ceramic Fig.5 Dental ceramic material systems for crown restoration [68]The zirconia ceramic restoration parts are commonly fabricated by slip casting or machining [61]. In recent years, machining based approach becomes increasingly adopted, in which the CAD/CAM technology is employed to produce ceramic parts with accurate shapes [27, 69]. In order to avoid introducing excessive surface damages to the ceramics, partially sintered ceramics or green ceramics can be machined, which will be subsequently sintered to achieve full mechanical strength. A process flow for this CAD/CAM process method is shown in Fig.6.

A process flow for this CAD/CAM process method is shown in Fig.6. After the scanning of the patients’ anatomy, the digital teeth models will be re-constructed, and a digital design of the dental restorations will be created. Based on the model, the CNC machining toolpath will be generated by the CAM software, and pre-sintered ceramic blocks will be machined to the shape of the restorations with up to 25% of scaling-up factor. Afterwards, the machined green ceramics will be sintered in a furnace to achieve the final mechanical strength. Due to the digitalization of dental restoration design procedures and the ease of machining with the green zirconia, this process route can readily achieve high production speed (~25-30min/part) and is reasonably cheap [7.8]. However, even with the green ceramic parts, micro-cracking is still likely to occur due to the loss of mechanical strength of exposed green part surfaces [25, 27].

However, even with the green ceramic parts, micro-cracking is still likely to occur due to the loss of mechanical strength of exposed green part surfaces [25, 27]. Although this could be alleviate by adding binder into the green part compacts or by pre-sintering of the parts, the micro-crack issue could not completely eliminated [25]. This creates potential concerns658with the long-term serviceability of the restoration parts and likely contributed partly to the higher failure rate of all-ceramic zirconia restorations.Fig.6 CAD/CAM for dental ceramic restoration manufacturingAnother relatively new ceramic dental material is the lithium disilicate (Li2Si2O5). Due it its relatively high degree of crystallization (~70% vol.), lithium disilicate possesses significantly higher mechanical strength and biocompatibility compared to the other glass ceramics [62, 64].

Lithium disilicate is usually made by precipitation from the SiO2-Li2O-K2O-ZnO-P2O5-Al2O3- La2O3 system, and the mismatch of the coefficient of thermal expansion between the lithium disilicate and the glass matrix is largely contributing to the improved fracture toughness of the material. With good clinical survival rate and high aesthetic quality, lithium disilicate possess potentials to be widely used in applications in inlays, onlays, crowns and other restorations. Fig.7 shows the comparison of a crown made by lithium disilicate and zirconia, and the difference in aesthetics is significant. Lithium disilicate dental parts are fabricated by either investment casting process or CAD/CAM process using the pre-made lithium disilicate feedstock, as in-process alloying is inadequate in developing the required crystal structure [71].a. Lithium disilicateb.

Lithium disilicateb. ZirconiaFig.7 Crown parts made by two ceramic materials [62]The use of AM in dental ceramic fabrication has been reported using several processes, although various issues remain unsolved [70, 72-75]. These issues including the lack of geometrical accuracy, staircase surface effects, and high porosities. For many dental prosthetic applications the marginal tolerance of the prostheses needs to be <0.1mm, which is difficult to achieve for many AM processes especially when material density and strength are also required [74]. In addition, due to the need of post-printing sintering processes with many AM processes, the porosity control and sintering distortion remain challenging. Compared to the CAD/CAM process, currently AM does not appear to have significant advantage with production speed, therefore for zirconia dental prostheses the justification might need to be established via performance enhancement in the future.

Furthermore, as previously mentioned, one of the unique requirements for dental ceramics is the need of specific optical characteristics. Although the fabrication of glass using AM with somewhat controlled optical properties has been demonstrated659[76-79], so far no technical ground appear to exist for the design of AM process or materials for tailored optical properties. As for the near future, such issues will likely have to be handled via either the use of specific materials (e.g. lithium dilisicate) with pre-defined optical characteristics or post-printing surface treatment processes (e.g. coating).ElectronicsThe use of ceramics in electronics applications is widespread. It was estimated that the global market for electronic ceramics was around $9 billion in 2014 [80]. In these applications, ceramic materials are used for their intrinsic dielectric properties and tailorable electrical and electromagnetic properties when doped.

In these applications, ceramic materials are used for their intrinsic dielectric properties and tailorable electrical and electromagnetic properties when doped. Ceramics are used as insulators, conductors, capacitors, piezoelectric sensors and actuators, semiconductors, optical and electromagnetic films and many other applications [81]. Therefore, the requirements of ceramic materials for this application category are most diverse. In many of these applications, a myriad of studies have elucidated the importance of microstructure including phases, grain sizes, grain boundary characteristics, contaminants/impurities, point defects (vacancy and interstitialcy) and phase morphologies, to name a few [82-87].

It is beneficial to design for small-size ceramic electronic components since this reduces the tendency of defect generation during the manufacturing processes [88], although the requirement of small size also comes from the demand of high performance efficiency and lightweight for many applications [89]. This implies that the manufacturing processes for these applications must address both the microstructural control and high-resolution geometry control while ensuring acceptable production rate, which is quite challenging for the current AM technologies. In addition, many of the ceramic electronics components and systems exhibit rather complex architecture as shown in Fig.8, which often involve the use of multiple materials that might cross categories (e.g. metals + ceramics).

metals + ceramics). As multi-material fabrication is still a relatively new area for AM, it might take considerable research efforts for the AM technology to become competitive in comparison to the traditional micro/nano fabrication technologies. On the other hand, due to the extensive employment of layerwise micro-fabrication processes such as photolithography, sputtering and vapor deposition, the current design architecture for these electronic devices could theoretically allow for smooth transition towards AM if the multi-scale manufacturing issues can be solved.

It could also be envisioned that for AM processes that are capable of voxel-based fabrication on non-flat surfaces such as the aerosol jet direct write technology [90], the capability to accommodate 3D integrated electronics design could potentially bring about disruptive application opportunities, as currently such research and development in such areas are significantly hindered by the existing fabrication processes despite the promising potentials [91-93].a. Solid oxide fuel cell [94]b. LED module package [95]c. Quantum dot solar cell [96]Fig.8 Architectures of some ceramic electronics660Current AM ceramics: state of the artAlmost all types of AM technologies except for material jetting have been employed for the fabrication of ceramic parts and structures. Most of the existing literature focus on the demonstration of AM ceramic fabrication capabilities with specific emphasis on geometry creations.

Most of the existing literature focus on the demonstration of AM ceramic fabrication capabilities with specific emphasis on geometry creations. In this paper, the emphasis was focused on the powder bed fusion, binder jetting, stereolithography and sheet lamination AM processes.Powder bed fusionPowder bed fusion (PBF) processes was among the first AM processes that were employed for ceramic fabrication. The PBF processes that utilize electron beam as a heat source have received little attention from the researchers for the fabrication of ceramic materials likely due to the thermally induced challenges such as cracking and structural distortion [97, 98]. On the other hand, the applicability of the PBF processes with a laser beam heating supply have been studied by different researchers. Silica sand components was fabricated by laser melting in an attempt to produce casting molds [99, 100].

Silica sand components was fabricated by laser melting in an attempt to produce casting molds [99, 100]. Fabrication yttria–zirconia powders via laser sintering/melting for dental applications was also experimentally investigated, although the resulting density of the structure was rather low (56%) [101].a. Part viewb. Top viewFig.15 Zirconia cube fabricated by laser sintering/melting [101]Yttria stabilized zirconia (YSZ) is among the few ceramic materials that have been investigated extensively using PBF processes, likely due to its relatively favorable phase stability after solidification and its high toughess. Shishkovsky et al. investigated the process of the mixture of alumina and ZST using laser melting PBF using different atmospheric conditions [102]. The mechanical performance of the fabricated ceramic parts was sensitive to the process parameters due to the formation of the porosities.

The mechanical performance of the fabricated ceramic parts was sensitive to the process parameters due to the formation of the porosities. Porosities of above 100 μm was observed, along with the occurrence of cracks as shown in Fig.15.1. The cracks were likely caused by the accumulation of heat during the process and is not common for laser melting PBF of this type of material and could be alleviated by preheating of the powder bed. Fig.15.2 shows a YSZ-Al2O3 dental bridge fabricated by laser melting, in which the powder was preheated to >1600°C [103]. The poor surface quality of the fabricated sample is attributed to the large melt pool size facilitated by high preheating temperatures as well as surface powder sintering. The authors reported that the fabricated parts were of almost 100% density with flexural strengths of above 500 MPa, and macroscale crack was largely prevented by the preheating.

The authors reported that the fabricated parts were of almost 100% density with flexural strengths of above 500 MPa, and macroscale crack was largely prevented by the preheating. On the other hand, due to the thermal shocks imposed during the spreading the cold ceramic powder, presence of fine cracks on the surface of fabricated components were observed.661a. Crack of YSZ-Al2O3 parts –in air (x50)b. Porosity – in argon (x20)Fig.15.1 Characteristics of the laser melted YSZ-Al2O3 [102]Fig.15.2 Dental bridge restoration printed via laser melting PBF [103]In general, for the powder bed fusion processes, fabrication of ceramic parts seems to be more challenging compared to other AM process, and most studies have focused to assess the feasibility of PBF technology for ceramic material systems.

With the successful demonstration of high quality fabrication of ceramic parts using the other AM technologies, the PBF does not draw extensive attention in this area.Binder jettingBinder jetting (BJ) has been extensively utilized for ceramic fabrication due to its unique process characteristics. Due to the use of binder as the geometry shaping mechanism, BJ process eliminates thermal inputs and could therefore process a broad range of materials including ceramics [104]. There exist an abundance of literature that demonstrated the use of BJ in the fabrication of various ceramics such as Al2O3, SiC, zirconia toughened alumina (ZTA), hydroxyapatite and calcium phosphate [105-122]. BJ has been investigated particularly for its potentials with bio-ceramics in biomedical applications [113-122]. Fig.16 shows two examples of the BJ ceramic structures for such applications.

BJ has been investigated particularly for its potentials with bio-ceramics in biomedical applications [113-122]. Fig.16 shows two examples of the BJ ceramic structures for such applications. This is largely because that in many biomedical applications porosity is required for the structures, which is one of the characteristics of powder based ceramic structures. Since the BJ process is essentially a green-part fabrication process, in many applications it can be used to readily substitute the traditional powder-based ceramic processes to directly fabricate porous ceramic structures. These porous ceramic structures are either used directly or subjected to further post-processing that can be guided by essentially the same guidelines developed for traditional powder-based green ceramic components.662a. Calcium phosphate scaffold [122]b.

Calcium phosphate scaffold [122]b. tri-calcium phosphate cranial segment [116]Fig.16 Bioceramic structures by BJOn the other hand, the BJ processes are subjected to similar constraints to the traditional powder based ceramic processes. Due to the powder spreadability requirement, the BJ processes often require that the powder feedstock exhibits good flowability [123]. This imposes significant constraints, as many ceramic powders are prepared via ball milling processes or chemical reaction based processes, and therefore often assume irregular particle morphology that exhibit low flowability and large aggregation [124]. Low flowability not only causes low packing density of the powder bed but also introduces significant spreading defects. In some works the ceramic powder was pre-treated with coating in the attempt to improve their flowaiblity. However, such approach inevitably introduces contamination phases that might be unwanted in applications [125, 126].

However, such approach inevitably introduces contamination phases that might be unwanted in applications [125, 126]. Another approach that might possess promising potential is the use of ceramic slurries as feedstock [127-130]. Such approach not only allows for the fabrication of low-flowability but also fine ceramic particles that are too small to be spread in dry status. On the other hand, since the green part printing resolution and accuracy of the BJ process is largely controlled by the binder liquid permeation process, the introduce of additional liquid phase will likely cause new issues with process control and optimization. In addition, manufacturing issues related to the use of ceramic slurry such as shrinkage crack and distortion must also be considered. Various post processing techniques have been employed to achieve good density for BJ green ceramic parts, such as liquid phase sintering, infiltration, and isostatic pressing [131-136].

Various post processing techniques have been employed to achieve good density for BJ green ceramic parts, such as liquid phase sintering, infiltration, and isostatic pressing [131-136]. The liquid phase sintering has been reported as an effective method in increasing the part density of BJ green ceramic parts [131-133]. Works have been reported for the infiltration of BJ alumina components by lanthanum-alumino-silicate glass and copper alloys, which resulted in the significant increase in the part density [134, 135], and in some cases up to 450% in flexural strength increase was observed in the specimens infiltrated with epoxy [136].

For the applications where the compositional purity of the fabricated samples is required, traditional ceramic processing technologies such as warm/cold isostatic pressing might be helpful in reaching high densities of components with relatively simple geometries [107, 137].StereolithographyA lot of the recent development efforts with AM ceramics focus on the use of stereolithography (SLA) based technologies. SLA is known for its high process resolution, and is therefore considered a promising technology for AM ceramic fabrication development considering the fact that post-processing of ceramics tend to introduce damages. Although the need for photopolymer feedstock poses significant constraints with the material development, various monolithic materials such as ZrO2, Al2O3, SiC, SiOC, SiO2 and TiC have been fabricated using SLA with two primary forms of material feedstock: (1) ceramic suspension, and (2) ceramic-663derivable polymer precursors.

There exist an abundance of literature about the development of ceramic suspensions for SLA. Such developmental work is not trivial, as the control of rheological and photocuring characteristics is of critical importance in the control of the SLA part qualities and more so the ceramic SLA [138-140]. For ceramic suspensions, these properties of the ceramic suspensions are influenced by various factors such as ceramic solid contents, dispersant and diluent concentrations, photoinitiator concentration and temperatures [141-145]. This is also closely related to the configuration of the SLA systems. For the conventional top-down configuration, in which the part is submerged in the photopolymer resin during the process, the resin is spread over uneven surfaces during the recoating. Therefore, the large variability of the liquid thickness imposes constraints to the development of the feedstock resin recipes [140].

Therefore, the large variability of the liquid thickness imposes constraints to the development of the feedstock resin recipes [140]. With bottom-up configuration, the part is placed upside-down with only the top layer of the part in contact with the resin, such recoating issue is alleviated, although the liquid phase separation and recoating upon each additional layer becomes an issue [146]. Lower viscosity is generally desired for the ceramic suspension [147], although this is often difficult to achieve due to the other considerations. To ensure effective sintering with the brown parts, it is suggested that a minimum solid content of 50% should be used for the ceramic suspensions [140, 148]. However, when the solid content exceeds 50%, the viscosity of the fluid increases drastically [149].

However, when the solid content exceeds 50%, the viscosity of the fluid increases drastically [149]. Solid contents of up to 60% has been reported, although porosities are still present in the final parts after sintering [141, 145, 150].Beside the rheological property issues, laser irradiation-related issues must also be considered for the development of ceramic suspension [140, 143, 144]. As shown in Fig.20, depending on the refractive index contrast, the resin feedstock could exhibit very different curing profiles under the UV irradiation due to scattering, which in turn influences the process planning and part qualities. In addition, the modulation of light intensity during different stages of the irradiation could also affect both the geometry and mechanical quality of the parts [151].

In addition, the modulation of light intensity during different stages of the irradiation could also affect both the geometry and mechanical quality of the parts [151]. Also, while excessive photocuring could reduce the part integrity and property, sufficient crosslinking is needed since the existence of residual monomer in the green ceramic parts could introduce cracking due to the thermally activated polymerization [152].Fig.20 Shape of the curing profile for SiO2 suspensions with varying refractive index values [140]Such dilemma between processibility and final part property for ceramic suspension based SLA may have prompted the development of the ceramic-derivable photopolymer as the feedstock. The principle of this approach is to utilize polymers that could decompose into ceramic products upon sintering/ pyrolysis for the green part fabrication.

The principle of this approach is to utilize polymers that could decompose into ceramic products upon sintering/ pyrolysis for the green part fabrication. Since the ceramic-derivable polymer precursors can be tailored to exhibit a broad range of electrical, optical, chemical and mechanical properties, the material development for SLA based ceramic systems appear to exhibit more potentials [153]. Additional compositional control could also be realized by adjusting the precursor recipes, which was demonstrated for the fabrication of SiC/SiOCN and SiC/SiOxCy composites664for engine turbine blades [60, 154]. In addition, since the pyrolysis temperatures of the polymer- derived-ceramics are relatively low (<1300°C), it also requires less energy for the sintering processes. Microfiber reinforced ceramic matrix has been fabricated, although the resulting structures still exhibit various defects such as surface striation and internal porosities.

Microfiber reinforced ceramic matrix has been fabricated, although the resulting structures still exhibit various defects such as surface striation and internal porosities. Recently, the use of microfiber and short fiber reinforcements in ceramic-derivable photopolymer was also demonstrated [60, 155]. The fabricated structures exhibit various defects such as surface striation and internal porosities, and the increased viscosity of the photopolymer due to the fiber addition could potentially become an issue. However, the increased viscosity could also be utilized to introduce shear-induced fiber alignment via micro-oscillation, as shown in Fig.

However, the increased viscosity could also be utilized to introduce shear-induced fiber alignment via micro-oscillation, as shown in Fig. 21 [155].Fig.21 Shear-induced fiber alignment in ceramic matrix composites, adapted from [155]Various literature has demonstrated the capability of SLA in the fabrication of complex ceramic geometries, such as the ceramic cellular structures [140, 153, 156] as shown in Fig.22a-b, turbine blades [60, 155] as shown in Fig.22c, investment casting cores [140] as shown in Fig.22d, dental crowns [74] as shown in Fig.22e and blood pump [157] as shown in Fig.22f. With vector and mask projection based SLA, minimum feature resolutions of 100-125µm were reported [150, 153, 158, 159]. For two-photo lithography based systems a minimum feature of 1.2µm was achieved for Al2O3, although the information about part quality is limited [160].a. SiOC cellular parts - green and final parts [153]b. SiO2 spring array [140]c. SiO2 blades [166]d.

SiOC cellular parts - green and final parts [153]b. SiO2 spring array [140]c. SiO2 blades [166]d. SiO2 investment core [140]e. Al2O3 dental crown [74] Fig.22 Ceramic geometries fabricated by SLAf. Ceramic blood pump [157]665On the other hand, porosity issue is still prominent with SLA ceramics, which often exhibit porosities ranging from 3-7%. The inevitable volumetric change resulted from the post-fabrication sintering process remains the main contributor to the existence of porosities. Some interesting attempts have been reported in remedy of this. In [161] the zirconia-toughened alumina (ZTA) was fabricated via stereolithography. The structure exhibit 99.5% of theoretical density and toughness value comparable to that of conventional processes. It was suggested that a careful two-step debinding process that combines vacuum pyrolysis and air debinding was largely responsible for the elimination of part defects [162].

It was suggested that a careful two-step debinding process that combines vacuum pyrolysis and air debinding was largely responsible for the elimination of part defects [162]. In another work by the same group, a liquid precursor infiltration process that serves to introduce additives or dopants to the printed ceramic parts was investigated [163]. The infiltration process was carried out after the debinding process, which improves the density and microstructure of the final parts after sintering. Such method was used on binder jetting processed samples as well [164]. The use of non-reactive plasticizing agent was also studied for the reduction of debinding defects, which reduces the amount of evaporated phase generated during the debinding [165].Sheet laminationSheet lamination (SL) utilizes sheets of materials as the feedstock for structure fabrication, which implies that the microstructure and quality of the material within a single layer can be well controlled.

The sheets used for part fabrication can be either cut and then stacked (Bond-then- Form) or stacked and then cut (Form-then-Bond) [166]. Although originally limited to paper feedstock, the SL principles have been successfully applied to the fabrication of metal, ceramic, and composite materials. In this case, ceramic tapes instead of paper sheets were used as build material for green part fabrication typically followed by high-temperature post-fabrication processing to debind and sinter the structures. Griffin et al fabricated ceramic components using alumina tapes via laminated object manufacturing (LOM) process, which exhibit comparable mechanical characteristics to pressed samples [167]. Further improvement in the bending strength of the fabricated alumina was accomplished by modifying the microstructure to an alternating composite structure of Al2O3 and Ce–ZrO2 layers from 311MPa for Al2O3 up to 688MPa for the alternating composite [168].

The use of binding agent is common for the SL with paper feedstock, however, when using a binding agent between successive ceramic tapes, unsuccessful debinding process might result in anisotropic mechanical properties in the fabricated components [167-169]. The applicability of SL technology for processing LiO2–ZrO2–SiO2–Al2O3 (LZSA) parent glass was investigated in various studies [170, 171], and 3D components have been successfully fabricated using LOM process as shown in Fig.23. However, high porosity values (greater than 10%) located at the lamination interfaces after post-processing limits the mechanical performance of the parts [170]. Other material systems such as glass-fiber reinforced polymers, SiC, SiC- fiber/SiC, AlN, Si3N4, HA, and TiC/Ni have also been examined as build material in sheet lamination technology [169, 172-175].

Due to the use of ceramic tapes, the SL processes usually only produce green parts, and consequent debinding and sintering must be accompanied by pressure in order to ensure the elimination of the porosities, and shrinkage along the layer stacking direction of less than 7% was observed due to the high green part densities [172]. The Reaction bonding Infiltration of LOM parts after sintering can be used to improve the density of the printed component [172]. The process flow of such approach is shown in Fig.24a. Successful demonstrate of SiC fiber reinforced SiC ceramic matrix composites was provided by the same group, which used alternating lay-up for consecutive ceramic layers to minimize fiber abrasion.

Successful demonstrate of SiC fiber reinforced SiC ceramic matrix composites was provided by the same group, which used alternating lay-up for consecutive ceramic layers to minimize fiber abrasion. However, it was observed that the interfaces between the SiC fibe and the SiC matrix were not well formed666(Fig.24b), which was caused by the lack of interfacial protection from oxidation during the sintering.The full potential of SL appears to be yet exploited, since this technology is among the few AM technologies that are readily capable of multi-material fabrication. The capability of the SL with non-flat surface also implies that the process might be applicable to the fabrication of complex ceramic laminates on the surfaces of existing structures [172], which might be a promising alternative for the fabrication of TBCs.

On the other hand, the difficulty in interfacial bonding quality control remain a challenge for SL to be employed for high-performance ceramic composites and TBCs. On the other hand, the z-direction resolution of the SL parts might be subjected to more constraints as the ceramic tape-pegs are required to possess minimum strength to be processed without damage. With the reported layer thickness of 0.25mm [173], the SL ceramic structures will be unsuitable for applications that require high geometrical accuracies such as dentistry.Fig.23 Gear wheel geometries: green laminate (left); sintered Li2O–ZrO2–SiO2–Al2O3 glass– ceramics (right) [171]b. Interfaces of SiC/SiC composite Fig.24 Fabrication of SiC/SiC fiber reinforced composite using SL + reaction bonding [172]a.

Interfaces of SiC/SiC composite Fig.24 Fabrication of SiC/SiC fiber reinforced composite using SL + reaction bonding [172]a. Process flow chartDiscussions and conclusionsIt appears that currently there exist various misalignments between the demonstrated AM ceramic capabilities and the needs for particular applications in the identified areas. While AM is generally capable of realizing relative complex geometries [176], such advantage is significantly compromised by the lack of microstructural quality control with the ceramic parts. Various issues such as porosity, purity, micro-defects and interfacial defects commonly exist with AM ceramic structures, which still require extensive efforts to overcome [137]. Due to the intrinsic staircase effect with the AM processes, the notch sensitivity issue of the ceramic parts is signified [27].

Due to the intrinsic staircase effect with the AM processes, the notch sensitivity issue of the ceramic parts is signified [27]. The need for post processes such as debinding and sintering implies that many of the challenges with the ceramic geometry designs that the traditional ceramic manufacturing face are also applicable667to AM ceramics, such as the limitation of feature thickness and the debinding/sintering distortion, which imposes geometry design limitations that greatly compromise the advantages offered by AM. In addition, from traditional powder based ceramic manufacturing it is known that finer ceramic particles are preferred for achieving higher densities during the sintering densification. However, refined particles tend to exhibit lower flowability and becomes difficult to spread using the powder bed AM approach.

However, refined particles tend to exhibit lower flowability and becomes difficult to spread using the powder bed AM approach. On the other hand, with the powder suspension feedstock based AM, due to the requirements with the rheological properties, the ceramic solid content is limited, which in turn limits the achievable density of the densified parts and necessitates more complex sintering strategies such as liquid phase sintering or HIP.Due to the need of post processes that are similar to the traditional ceramic manufacturing processes, AM could considerably benefit from the existing expertise within ceramic industries. For example, for slurry and suspension based ceramic AM technologies, there exist extensive knowledge base in this subject that has yet to be explored within the context of AM. For example, various literatures have provided excellent reviews about the key consideration factors for the development of the ceramic pastes during the extrusion based shaping processes.

For example, various literatures have provided excellent reviews about the key consideration factors for the development of the ceramic pastes during the extrusion based shaping processes. The flow rheology of the viscous ceramic pastes can be modeled, and phase separation at macro-scale must be avoided as it could lead to defects such as cracks [177]. During the material extrusion processes, flow-induced alignment of the ceramic particles could cause additional anisotropy in the structures, which can be potentially utilized to create textured microstructure [178]. The paste must be free of large aggregates in order to produce defect-free structures.

The paste must be free of large aggregates in order to produce defect-free structures. In addition, during the debinding of the green ceramic parts made of powder injection molding, broader binder decomposition temperature ranges, which can be characterized by either thermogravimetric analysis or digital scanning calorimetry, benefit the reduction of the crack formation in the de-binded brown parts, since the escape process of the evaporated gas phase from the binder is more gradual [178]. Slower binder decomposition is also beneficial for the retaining of the part shape. These knowledge will likely facilitate the development of new binder and slurry systems for the AM ceramics, which is an area that does not appear to be explored yet.In terms of ceramic material availability, the existing AM technologies already demonstrated excellent material compatibility.

New materials such as boron carbide (B4C) and titanium boride (TiB2) might also possess promising potentials for AM adoption due to their potentials in high value-added ceramic armor applications [179, 180]. In addition, the capability of AM in applying materials on a voxel-by-voxel basis could be potentially utilized for the doping of ceramics, which is an important method in altering the material characteristics of ceramics [181].In the near term, in applications where functional ceramic structures are needed that use material feedstock already exhibiting the intended performance, the geometry realization potential of the AM could be better utilized [182]. For example, the fabrication of piezoelectric ceramic devices using high-piezoelectric coefficient ceramics could yield parts that are comparable to the traditional ones [182]. In addition, in specific applications where exotic geometries such as cellular structures are desired, AM could also find immediate use.

In addition, in specific applications where exotic geometries such as cellular structures are desired, AM could also find immediate use. For example, it was claimed that the printed HA implants have the required porosity for osseointegration while achieving mechanical strength 3-5 times better compared to the conventional implants with the same material and porosity [183]. In the long term, ceramic AM will likely benefit from close collaborative research and development efforts between the academia, system manufacturers and industrial users, as668application-oriented development works with the platform, the materials and the process optimizations are all needed.AcknowledgementThe authors would like to thank Dr. Johannes Homa and Shawn Allan for the helpfuldiscussions about the subject.References [1] S. Somiya. Advanced Technical Ceramics. Academic Press Inc, San Diego, CA, USA, 1989. [2] R. Naslain.

ramic Joining by Gas Phase Pulsed Laser Processing C. M. Weiss, M. Aindow, H. Marcus Institute of Materials Science, Materials Science and Engineering Program/CMBE, University of Connecticut, StorrsAbstractThe method of Selective Area Laser Deposition (SALD) and Vapor Infiltration (SALDVI) has been successfully used to fabricate small three-dimensional SiC/SiC and SiC/metal powder parts. Ceramic joints made by this technique have been limited by the throwing power of the laser resulting in incomplete joint penetration. Studies were performed to show the effectiveness of a fiber laser, with a wavelength of 1070 nm, for a joining process. The ability of the laser to penetrate a powder bed was utilized in the joint fabrication. The combination of powder fill, and deep laser penetration into the powder bed shows potential in the field of ceramic joining.IntroductionCeramic joining of components can be done by several methods.

Green pieces can be joined ultrasonically before firing, frictional joining or the use of brazing materials can be used if the components are already fabricated. These methods involve special conditions that must be considered before component design. Another method that has been studied is the use of a focused laser in a reactive gas environment to selectively heat a substrate and decompose the ceramic precursor and form a joint.[3,6-9] This process, known as selective area laser deposition (SALD), can be performed without bulk heating of the components and without the use of materials that will be ineffective at higher temperatures. The SALD has been used with some success to join components with a rotational symmetry.[7] A variant of the SALD process that involves the infiltration of a substrate powder with precursor gas during decomposition is known as the SALDVI process where the VI indicates vapor infiltration.

This technique has been used for free form fabrication of arbitrarily shaped components. The powder that is to be infiltrated can be the same material as the decomposed species or a different material making a composite. Several types of ceramics have been fabricated with the SALD and SALDVI process including Si3N4, graphite and composite materials utilizing different metallic and ceramic powder reinforcements. This study utilizes Tetramethylsilane (Si(CH3)4) pyrolytic decomposition in the presence of hydrogen in the overall process reaction: Si(CH3)4(g) ⇒ SiC + 3CH4(g) Alternative precursors to form SiC include methyltrichlorosilane (MTS , SiCH3Cl3), and combination of silane (SiH4) and methane (CH4). [5,9-13]Experimental ProcedurePrevious joining work has concentrated on the use of deposition without the use of a powder to infiltrate. This study combines the use of SALDVI with the previously established joining techniques.

This study combines the use of SALDVI with the previously established joining techniques. A 40-watt average power pulsed fiber laser with 1.5 mJ per pulse (SPI Lasers) was used for the studies. A single focusing lens was used with a focused spot of 50 microns. The beam was delivered off of a series of mirrors on an XY motion stage controlled by a Labview motion program and a National Instruments control card. The deposition chamber is a vacuum chamber flushed with argon and evacuated to 80 mtorr and then filled with desired gas proportions. The starting gas operating pressure for all experiments was 50 torr TMS and 50 torrH2,below the vapor pressure of TMS, to prevent condensation within the chamber. Semi Infinite powder bed deposition was used to evaluate deposition conditions using a 7 micron average size SiC powder. Laser powers of 20 and 38 watts average were used with laser scan speeds of 100, 200, and 400 microns/s.

Laser powers of 20 and 38 watts average were used with laser scan speeds of 100, 200, and 400 microns/s. The higher power at all speeds ejected considerable powder from the point of focus leaving trenches in the powder with deposits into the trench walls. With the lower power the most consistent deposit was produced at 100 micron/s. To determine the spacing to be used for the joint deposition raster pattern, powder bed deposits were made with decreasing line spacing from 1800 microns to 80 microns. A powder bed raster deposit was made 1.5 mm wide with a scan line spacing of 85 microns and a total length of 6 mm and a scan speed of 100 micron/s at 20 watts average power.The joint configuration explored was an abutting of two beveled edges with a powder fill of 7 micron SiC powder, packed into the joint cavity by ultrasonic vibration. A joint component thickness was chosen so that the depth of penetration of the laser into a free powder bed was equivalent to the SiC plate.

A joint component thickness was chosen so that the depth of penetration of the laser into a free powder bed was equivalent to the SiC plate. A parallel rectangular raster pattern was chosen with a line spacing of 130 microns in the travel direction. The deposited width of the raster pattern was designed as 130% of the largest part of the joint gap. The sample geometry used in this study is indicated below, Figure 1.dacbFigure 1: Schematic of sample configuration: a. Beveled SiC plates 1270 micron thick with standardized 20° bevels. b. Diffusion gap created by spacers to allow diffusion of reactant as well as a measure so the joint would not adhere to the sample mount. c. Mounting clamp that secures each beveled section. d. Rectangular raster beam path.The two beveled SiC plates were abutted and clamped onto a sample stage.

c. Mounting clamp that secures each beveled section. d. Rectangular raster beam path.The two beveled SiC plates were abutted and clamped onto a sample stage. The samples were separated from the sample stage by thin sheet steel spacers to prevent any adhesion to the stage and to allow for possible diffusion of the reactant species to the non-beveled side of the joint. Two variations of the joint were evaluated, both were formed with a 50 micron spot size scanned raster with a 20 watt average power with a starting TMS pressure of 50 torr, H2 pressure of 50 torr, and a scan speed of 100 micron/s. The first joint fabrication type was performed with no joint surface preparation. The second joint was formed by pre ablating the faces of the joint components without the presence of powder or reactive gas in order to produce a micro textured surface and then filling the joint cavity with powder and performing the SALDVI process as had been performed in the joint type 1.

Analysis of the deposited material was performed with the use of visible light microscopy, environmental scanning electron microscopy as well as X-ray diffraction.Results and DiscussionThe decomposition of TMS in this study has two general microstructural types. The SiC formed by the infiltration of the powder and the SiC that has formed either by displacing powder or by decomposing on already infiltrated powder. The powder infiltration deposit type is seen below, Figure 2.a Figure 2: The decomposition product has surrounded the SiC grains with a somewhat dense deposit. a. Infiltrated powder at 25X visible light magnification taken from a progressive line spacing scan in a region of uniform deposition. b. a 125X visible light image taken from an infiltrated region of a joint sample. c.

b. a 125X visible light image taken from an infiltrated region of a joint sample. c. an ESEM image of a powder bed scanned deposit at 2500X magnificationbcWhen infiltration paths into the powder bed are closed by the growth of the SiC deposit, acondition arises that allows SALD to occur, deposition without infiltration, Figure 3.abFigure 3: These images are representative of the decompositions that occur without powder infiltration. a. an ESEM image at a magnification of 405X of the deposit above the powder bed. Image b is a higher magnification, 1000X, of the same region above the powder. Reasonable densities are observed.To determine the optimal spacing for the parallel lines of the raster pattern a deposition was performed with variable line spacing. The factors considered for optimal spacing were: uniformity of the deposit in the vicinity of a scanned line, the extent to which an adjacent line was resolvable and the presence of beam damage.

The factors considered for optimal spacing were: uniformity of the deposit in the vicinity of a scanned line, the extent to which an adjacent line was resolvable and the presence of beam damage. Light microscopy was used to evaluate the beam spacing, Figure 4.abFigure 4: The proximity of one line to the next in the powder bed deposition, affects the quality of the deposit. a. the central line is epoxy material which has filled the space left by the laser powder displacement and this micrograph represents the most widely spaced scan width (1.8mm). b. an intermediary scan spacing (300 microns).The spacings of the raster lines that produced the most uniform deposits were found to be in the range of 300-80 micron. Deposition into a powder bed with a raster pattern with line spacing of 85 microns was performed on a semi-infinite powder bed. Resulting in a cohesive unit of good uniformity and reasonable densityDeposition for the purpose of joint formation was performed.

Resulting in a cohesive unit of good uniformity and reasonable densityDeposition for the purpose of joint formation was performed. The goal of the joining studies is the formation of a strong and minimally porous joint with a full penetration to the base of the joint, with good cohesion to the joint wall. Two joining techniques were attempted. Joint type 1 is characterized by flat joint wall surface prepared by diamond grinding (to 220 grit). Joint type 2 was laser scribed by the fiber laser in order to provide a surface roughness. Previous joining attempts were met with greater success when the joining surfaces showed some porosity.[8]The porosity of each joint was qualitatively evaluated by the ESEM. The initial particle size for the substrate powder is less than 7 microns. The structure of the deposition indicates the growth of SiC surrounding the particles, Figure 5.abFigure 5 a.

The initial particle size for the substrate powder is less than 7 microns. The structure of the deposition indicates the growth of SiC surrounding the particles, Figure 5.abFigure 5 a. Deposited joint material with a 1000 X magnification that displays a region of partially reacted powder in the upper left next to a larger region of deposited material. It can be seen that deposition was occurring with some irregularly in this sample, joint type 1. Image b taken from a sample of joint type 2 at a magnification of 2500 X displays a more uniform deposition with no regions of uncoated powder. Both images display reasonable density.The use of a high quality beam source, in the form of a fiber laser, and a focusing optic with a 50-micron spot size, has allowed a deep penetration into the joint cavity. Although both joints experienced a de-cohesion as a result of moments applied during the sample removal, the joint type 2 shows evidence of deep joint penetration.

Although both joints experienced a de-cohesion as a result of moments applied during the sample removal, the joint type 2 shows evidence of deep joint penetration. Although fracture occurred along one deposition edge, the un-fractured edge shows little porosity at the boundary, Figure 6.abcFigure 6 a. A polished sample taken from a joint type 1 on an optical microscope with a magnification of 50X that shows a distinct and non porous boundary between the deposit material, I, and the joint wall/substrate, II. ESEM images of Joints type 1 (b.), and Joint type 2 (c.) where both show the same type of adhesion at a magnification of 500X where the left side of each interface is the deposit. c. shows a region, identified by an arrow, where the deposition material penetrated a micro ablation present from the laser pretreatment.The beam penetration is evident from the deposited material in the presence of the crack tip.

The joint type 1 joint did not have full penetration to the joint tip. There is evidence of deposited material at the tip of the second joint type, Figure 7.abcdFigure 7 a. Micrograph of the type 1 joint indicating partial penetration before joint fracture, image at a magnification of 125 X(backscattered ESEM). b. Optical micrograph (12.5 X magnification) of the type 2 joint penetration, the deposited material (black) is adhered to the joint tip. The grey material to the left of the circled region is infiltrated epoxy into the fracture zone. Image c is an optical magnification of the same region at 50X magnification. Image d is an ESEM image with a magnification of 1000X. Image d also shows evidence of the laser pre ablation in the substrate tip, right.Deposits were confirmed to be silicon carbide through the use of powder x-ray diffraction.

Image d also shows evidence of the laser pre ablation in the substrate tip, right.Deposits were confirmed to be silicon carbide through the use of powder x-ray diffraction. The diffraction pattern standard for synthetic hexagonal silicon carbide corresponds closely to the experimental pattern as seen below, Figure 8.Figure 8 X-Ray Diffraction pattern for the composite joint material compared with a standard for hexagonal SiC.ConclusionThe novel technologies employed in this study include the use of a high quality fiber laser generated beam focused to 50 micron, the use of powder material filler phase within the joint, and the use of the laser as a means of ablating the surface of the joint before the infiltration process. The powder bed infiltration tests indicate that the beam is capable of penetrating the powder and decomposing the precursor to a uniform depth. If the sample joint is kept to a thickness below that of the penetration depth a SALDVI type joint is possible.

If the sample joint is kept to a thickness below that of the penetration depth a SALDVI type joint is possible. Thicker samples would necessitate a multi layer approach. The fiber beam shows promise for future joining studies but a balance must be reached between the desired penetration depth of the beam and the tendency for the beam to displace the powder from the hot zone. The surface pre ablation technique employed here was detectable under microscopic examination. There was deposited material within the ablation tracks. The ablation pattern was made under potentially oxidizing conditions, which could have been a factor in joint de-cohesion. Ongoing work involves the use of an ablation pattern applied in a vacuum before powder infiltration. Additionally a multi layer powder infiltration joining technique is being explored.References[1] B. R. Birmingham, J. V. Tompkins, G. Zong, H. L.

ramic Joining by Selective Beam DepositionJ. V. Tompkins, B. R. Birmingham, H. L. Marcus Center for Materials Science and Engineering The University of Texas at Austin Austin, TXAbstractCurrent methods of joining of ceramic components may compromise the strength, chemical resistance, or high temperature properties of the resulting ceramic parts. A new method of joining, Ceramic Joining by Selective Beam Deposition, creates an all-ceramic joint between two or more ceramic components through selective decomposition of a gas precursor. An all-ceramic joint not only preserves the valuable properties of the ceramic materials joined, but may be tailored to match the coefficient of thermal expansion of the original material(s). The added material may be the same as one or both of the joined Inaterials, or may be a composite material.

The added material may be the same as one or both of the joined Inaterials, or may be a composite material. This preliminary work explores the effect of scanning speed and precursor pressure on Selective Beam Deposition of silicon carbide using tetramethylsilane.IntroductionThe need to join ceramic components has driven the development of techniques such as brazing with metals or low temperature binders, ultrasonic joining of green pieces prior to firing, or friction welding. The following describes a new process, Ceramic Joining by Selective Beam Deposition, which involves placing ceramic objects together in a desired configuration within a reactive gas environment and scanning a laser beam over the contact area of the parts. The laser beam causes a gas reaction and produces new ceramic material in the area to be joined.

The laser beam causes a gas reaction and produces new ceramic material in the area to be joined. Since gas reactions can produce advanced ceramics, the process eliminates the need for low temperature binder phases which often have low resistance to temperature, chemical corrosion, or abrasion. An all ceramic joint may also avoid problems with thermal expansion as well as wetting in the interface. Since gas reactions may occur well below melting or sintering temperatures the process minimizes thermal stress. It also works on ceramics which decompose below their melting point, such as silicon nitride. Additionally, the process can join materials with very different melting temperatures (for instance, a ceramic and a metal).

Additionally, the process can join materials with very different melting temperatures (for instance, a ceramic and a metal). Control of gas compositions allows deposition of composite materials within the joint, either to wet two different phases more effectively or to mitigate thermal expansion mismatch between joined materials.The following work uses tetramethylsilane (Si(CH3)4) as the precursor gas. Pyrolysis of tetramethylsilane (TMS) occurs according to the following overall reaction [1]:Si(CH3)4 (g) ==> SiC + 3 CH4 (g)Previous work in Selective Area Laser Deposition (SALD) has demonstrated that deposition rates up to 6 mm/min may occur where a focused laser beam strikes a substrate surface. Recent SALD work has included deposition of pyrolytic carbon [2,3], silicon carbide, silicon nitride [4,5], titanium [6], and nickel [7].

Recent SALD work has included deposition of pyrolytic carbon [2,3], silicon carbide, silicon nitride [4,5], titanium [6], and nickel [7]. In the case of silicon carbide, growth rates and hardness values indicated that selective deposition offers a viable method of joining advanced ceramic materials.409ProcedureThe equipment used for evaluating the joining process is similar to that used for SALD [8] with Fig. 1 illustrating specific differences between the two. A rotating shaft holds sections of ceramic tube together within a vacuum chamber containing TMS vapor at the desired pressure. A focused beam from a laser strikes the adjacent surfaces of the tube sections, decomposing the TMS to produce silicon carbide in the joint. A flask containing liquid TMS provides vapor at pressures up to 600 Torr at room temperature. Auxiliary gas lines provide H2 and N2 as necessary.

A flask containing liquid TMS provides vapor at pressures up to 600 Torr at room temperature. Auxiliary gas lines provide H2 and N2 as necessary. To provide forced convection at the deposition point, the chamber contains a small fan.The chemical and thermal processes occurring at the joint during deposition resemble those shown in Fig. 2. TMS diffuses to the heated area at the surface where it decomposes to form silicon carbide and methane. At the same time, methane diffuses away from the surface. A bevel in the ceramic tubes at the contact area increases the area available for adhesion and allows a flush surface after the addition of silicon carbide.Previous SALD work has shown that the growth rate and mechanical properties of deposited silicon carbide depend on variables such gas composition, scan rate, and laser power. Growth rate increases in a roughly linear fashion with TMS pressure [9].

Growth rate increases in a roughly linear fashion with TMS pressure [9]. With this in mind, the following work used TMS pressures of up to 350 Torr, as pressures over 350 Torr cause excessive formation of a solid by-product on the walls of the vacuum chamber. Earlier work with Selective Area Laser Deposition also revealed that scan rate affects the morphology of deposited material [10]. Finally, in the past growth rate has increased with increasing laser power. Actual process temperature measurements require a non-contact method which can distinguish a circular region approximately 0.5 mm in diameter, which the current system does not include. Consequently, we cite laser power in Watts and the calculated Ve 2 spot size in lieu of surface temperature at the deposition point.Ceramic Joining by Selective Beam Deposition introduces parameters in addition to those listed above for SALD. Chief among these are the physical properties of the materials to be joined.

Chief among these are the physical properties of the materials to be joined. The reflectivity and absorptivity of the joined materials in addition to those of the material(s) being deposited dictate the coupling of the laser beam to the deposition surface, and hence, the beam power required. Since these properties vary with the wavelength of laser light used, the materials selected may determine the type of laser which is appropriate. In this work, we considered alumina and silicon carbide objects for joining with deposited silicon carbide. Alumina possesses low intrinsic absorptivity for both the C02 and Nd-Yag lasers available, while silicon carbide showed high reflectivity for the C02 laser in preliminary trials. Both materials also have different thermal conductivities. To find a suitable rnaterial/laser combination, both materials were tested for deposition in a fixed spot as well as in a moving beam with both lasers.

To find a suitable rnaterial/laser combination, both materials were tested for deposition in a fixed spot as well as in a moving beam with both lasers. Also, the porosity of the material to be joined affects both the optical properties and the thermal conductivity. Therefore, substrate porosity was included as a variable in the above tests.Another variable introduced by the joining aspect of the process is the geometry of the edges of the material to be joined. A simple 45° bevel was selected as a standard geometry for all trials. However, it is anticipated that varying bevel geometry will have an effect on beam absorption, heat conduction, mass transport, and adhesion of the deposited material.After determining suitable conditions for joining experiments, the next goal was to examine the effect of TMS pressure and scan rate on the process while holding other variables constant.

The Selective Beam Deposition process involves many transport, kinetic, and thermodynamic phenomena which make predicting the effect of deposition410Laser Beam InManometerWindowTMS, N2 inCeramic Tube SectionsStepper MotorVacuum LineFig 1. Schematic view of equipment used to evaluate Ceramic Joining by Selected Beam Deposition.Laser BeamI Methane DiffusionTMS Diffusion~~ I ~ I Parts to ~ " ' " S'l' be joinedtb'd 1 ICon car 1 eFig. 2 Detail of deposition area during joining process.411In view of this, the experiment spanned a wide range of both gas conditions very difficult. composition and scan rate to find trends which might guide further experiments. An 18mm diameter alumina tube was placed in the chamber with an initial TMS pressure of 350 Torr. A 12 Watt beam from a Nd-Yag laser was focused to a 0.5 mm spot on the tube, and a stepper motor rotated the tube at 2 RPH for a period of 15 hours.

A 12 Watt beam from a Nd-Yag laser was focused to a 0.5 mm spot on the tube, and a stepper motor rotated the tube at 2 RPH for a period of 15 hours. As material formed on the tube, the effective diameter of the tube increased, leading to a proportional increase in scan speed as the deposit thickness increased. Simultaneously, a data acquisition system recorded the change in total pressure inside the chamber as the deposition progressed. The resulting structure after a number of hours would be a disk with many layers grown at different conditions. Each layer can be observed for growth rate and uniformity, and if desired, hardness. The radial position of a given layer indicates the scan rate and gas pressure during the growth of that material.

Each layer can be observed for growth rate and uniformity, and if desired, hardness. The radial position of a given layer indicates the scan rate and gas pressure during the growth of that material. Although the effects of gas composition and scan rate could not be evaluated independently, this technique provided an easy method to view a wide range of deposition conditions to evaluate deposition trends.ResultsExperiments performed with the two ceramics at various porosities with the two available lasers indicated that neither the C02 laser (30W) nor the Nd-Yag laser (150W) caused deposition on fully dense silicon carbide or alumina. Alumina with a porosity of 20% or greater absorbed sufficient heat from the Nd-Yag beam to cause deposition of silicon carbide, but did not work with the lower power C02 beam. Silicon carbide with 50% porosity worked with both lasers. With these results, 80% dense alumina tubes were selected for joining experiments with the Nd-Yag beam as shown in Fig.

Silicon carbide with 50% porosity worked with both lasers. With these results, 80% dense alumina tubes were selected for joining experiments with the Nd-Yag beam as shown in Fig. 1 above. However, attempts to join beveled tubes at gas pressures of 50 and 350 Torr did not produce integral parts. At 50 Torr, the joint showed good adhesion and uniformity, but suffered from a lack of deposited material in a longitudinal line down the center of the joint. At 350 Torr, the deposited material also showed good adhesion, but grew in an uneven, rod-like fashion. The lack of deposited material at lower pressures resembles a type of growth previously observed in SALD work, and may be due to depletion of the reactant in the center of the deposition.The final experiment evaluated a wide range of pressures and scan rates to find a case where reactant concentration was sufficient to avoid depletion in the center of the deposit yet not high enough to produce rods or projections.

The result was a disc of uniform thickness of approximately 2 mm with clearly distinguishable layers as shown in Figs. 3a and 3b. The initial layers show evidence of non-uniform, rod-like growth. Following layers become more uniform as the deposition time increased. TMS partial pressure is estimated by assuming the total pressure consists of unreacted TMS and the methane byproduct from the overall reaction Si(CH3)4 (g) ~ SiC + 3 CH4 (g). Layer thickness vs. radial position was measured for three different radial sections. The decreasing thickness of layers over time initially appeared to result from both the decreased residence time as the radius increases and the decrease in reactant concentration. However, Fig. 4 shows the growth rate as a function of TMS pressure by dividing layer thickness by the residence time at a particular radius. The highest rates actually occurred several hours into the experiment after considerable depletion of TMS.

The highest rates actually occurred several hours into the experiment after considerable depletion of TMS. Aside from the information available from the material produced at various conditions in the disk itself, the initial layers in the disk adhered to both ceramic tube sections, successfully bonding them together.41298•l!iI!l!iI!IIIIIIII IIII •II IIIIl!iI!••IIIIIIIIIIIIIIl!iI!II 1111IIIIIIIIl!iI!IIIIl!iI!l!iI!l!iI!l!iI!IIIIIIII II.11IIIIo5101520Radial Position (mm)Fig. 4 Plot of growth rate vs. radial position. (Combined data from three different radii)DiscussionControl of deposition uniformity is the current limitation in the ability to perform Ceramic Joining by Selective Beam Deposition. Deposition rates for silicon carbide are already sufficient, and the material adheres strongly to substrates with 20% or more porosity.

Deposition rates for silicon carbide are already sufficient, and the material adheres strongly to substrates with 20% or more porosity. Uniformity suffers from an apparent starvation of reactant in the center of the deposition area at 50 Torr, and non-uniform growth at 300 Torr. The intermediate range may include conditions which avoid both problems. The disk produced by an extended deposition time investigates only certain combinations of scan rate and tetramethylsilane pressure. However, the ability to identify and examine material grown over a wide range of conditions provides a great deal of information in a single experiment. A continuously flowing gas reactor would avoid the accumulation of methane gas which occurs in the present system, but as a first order approximation we assume the effect of decreasing the activity of TMS on the formation of silicon carbide is independent of the presence of methane gas. Fig.

Fig. 4 gives a summary of the phenomena observed: at the beginning of the experiment, the growth rate increases gradually even though the TMS concentration decreases and the methane concentration increases. After approximately eight hours, the growth rate peaked. The gas composition, assuming the overall reaction shown above, was 75 Torr TMS and 675 Torr methane. Layer uniformity continued to improve as the experiment progressed.The increase in growth rate during a simultaneous decrease in TMS pressure and increase in methane pressure requires further study. One possible explanation is that the methane gas is producing pyrolytic carbon which is added to the silicon carbide growth. This possibility was not expected based on the lack of growth of pyrolitic carbon in earlier SALD experiments using methane [11]. It is possible that the presence of TMS acts to414II25facilitate the growth of carbon.

It is possible that the presence of TMS acts to414II25facilitate the growth of carbon. This explanation may be tested by examining the deposited material to establish the presence of a large fraction of pyrolitic carbon. If the increased growth rate is actually due to scan speed, scan speed may represent a useful parameter in decreasing the time required to join components, while at the same time increasing the surface uniformity of joint. A continuous flow reactor would help determine if this is an actual benefit by holding the gas composition constant as the radius of the disk increased, increasing the scan rate. Finally, one must consider that a phenomena not yet identified is responsible for the apparent increase in growth rate with respect to decreasing reactant concentration.The tendency for the process to produce distinct layers indicates that interfaces between successive scans must be addressed to produces a structurally sound joints.

In the past, the addition of hydrogen has improved interfaces in deposited silicon carbide [12].ConclusionThe above suggests important implications for a gas-based joining process. In general, the process will require high growth rates and good uniformity. The physical properties of the materials joined impose additional constraints to those seen in earlier SALD processes. While investigating a wide range of deposition conditions in an extended experiment, two ceramic objects were successfully joined by the process for the first time.AcknowledgementsThis material is based in part upon work supported by the Texas Advanced Technology Program under grant No. CSME-ATPD-157 and the Office of Naval Research grant No. N00014-92-J-1514.References1. Figueras, A., S. Garelik, J. Santiso, R. Rodroguez-Clemente, B. Armas, Combescure, R. Berjoan, J.M. Saurel, R.

N00014-92-J-1514.References1. Figueras, A., S. Garelik, J. Santiso, R. Rodroguez-Clemente, B. Armas, Combescure, R. Berjoan, J.M. Saurel, R. Caplain, "Growth and Properties of CVD-SiC Layers Using Tetramethylsilane", Materials Science and Engineering, Volume B11, No. 1 4, January 15, 1992, pp 83-88. 2. "Processing Problems Associated With Gas Phase Solid Freeform Fabrication Using Pyrolytic Selective Area Laser Deposition", Proceedings ofthe Solid Freeform Fabrication Symposium, edited by Harris L. Marcus, Joeseph J. Beaman, David L. Bourell, and Richard H. Crawford, The University of Texas at Austin, Austin, Texas, August 12-14, 1991, P 271. 3. Tapered Microstructures", Proceedings of the Solid Freeform Fabrication Symposium, edited by HarFis L. Marcus, J. J. Beaman, Joel W. Barlow, David L. Bourell, and Richard H. Crawford, ISSN 1053-2153, The University of Texas at Austin, Austin, Texas, August 7-9, 1995. 4.

Marcus, J. J. Beaman, Joel W. Barlow, David L. Bourell, and Richard H. Crawford, ISSN 1053-2153, The University of Texas at Austin, Austin, Texas, August 7-9, 1995. 4. Area Laser Deposition of Silicon Carbide", Proceedings of the Solid Freeform Fabrication Symposium, edited by Harris L. Marcus, J. J. Beaman, Joel W. Barlow, David Bourell, and Richard H. Crawford, ISSN 1053-2153, The University of Texas at Austin, Austin, Texas, August, 1994. 5. Wallenberger, F. T., "Amorphous Silicon NitrideFibers Grown from the Vapor Phase", Journal of Materials Research, v 9, n 3, March, 1994, pp 527-530. 415Zong, G., J. V. Tompkins, W. R. Thissell,Sajot, and H. L. Marcus,Pegna, Joeseph and James Maxwell, "Gas Phase Laser-Induced Pyrolysis of6. Proceedings of the Solid Freeform Fabrication Symposium, edited by Harris L. Marcus, J. J. Beaman, Joel W. Barlow, David L. Bourell, and Richard H. Crawford, ISSN 1053 2153, The University of Texas at Austin, Austin, Texas, August 7-9, 1995. 7.

Marcus, J. J. Beaman, Joel W. Barlow, David L. Bourell, and Richard H. Crawford, ISSN 1053 2153, The University of Texas at Austin, Austin, Texas, August 7-9, 1995. 7. Tapered Microstructures", Proceedings of the Solid Freeform Fabrication Symposium, edited by Harris L. Marcus, J. J. Beaman, Joel W. Barlow, David L. Bourell, and Richard H. Crawford, ISSN 1053-2153, The University of Texas at Austin, Austin, Texas, August 7-9, 1995. 8. Area Laser Deposition of Silicon Carbide", Proceedings ofthe Solid Freeform Fabrication Symposium, edited by Harris L. Marcus, J. J. Beaman, Joel W. Barlow, David Bourell, and Richard H. Crawford, ISSN 1053-2153, The University of Texas at Austin, Austin, Texas, August, 1994. 9. ibid 10. ibid 11. "Processing Problems Associated With Gas Phase Solid Freeform Fabrication Using Pyrolytic Selective Area Laser Deposition", Proceedings ofthe Solid Freeform Fabrication Symposium, edited by Harris L. Marcus, Joeseph J. Beaman, David L. Bourell, and Richard H.

viewed PaperCeramic On-Demand Extrusion of ZrB2-SiC Microchannels for Ultra High-Temperature Compact Heat ExchangerAbid Hasan Rafi1, Marharyta Lakusta2, Vincenzo Anthony1, Grace Basler2, Sophie Grier2, Lauren Moeller2, David Lipke2, Jeremy Lee Watts2, Greg Hilmas2, Ming C Leu11 Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, Missouri2 Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MissouriAbstractCeramic On-Demand Extrusion (CODE) is a paste extrusion based additive manufacturing process that has been developed for the fabrication of dense ceramic components. This paper presents a study of using the CODE process to build microchannels with ZrB2-SiC aqueous paste for the fabrication of an ultra-high-temperature compact heat exchanger (CHE).

This paper presents a study of using the CODE process to build microchannels with ZrB2-SiC aqueous paste for the fabrication of an ultra-high-temperature compact heat exchanger (CHE). The aqueous paste is developed from ZrB2 and SiC (70/30 vol%) powders to possess properties suitable for the extrusion process. The performance and properties of the microchannels are tested for use in the CHE, which features a large heat transfer area to volume ratio and has potential applications in aero-engines, microturbines, etc., as well as high-efficiency power generation systems.Keywords: Additive Manufacturing, Ceramic On-Demand Extrusion, High-temperature Ceramic, Compact Heat Exchanger, Microchannel, Paste Extrusion1.

IntroductionThe supercritical carbon dioxide (sCO2) Brayton cycle has gotten a lot of interest in recent years for use in energy applications like waste heat recovery [1,2], solar-thermal systems [3-4], fuel cells [5,6], coal power [7-9], geothermal power [10], and nuclear power [11,12] because of its high efficiency, small footprint, and simple layout. sCO2 has a higher diffusivity, lower viscosity, and lower surface tension than liquid CO2. For similar turbine intake temperatures, several studies have found that sCO2 Brayton power cycles have higher cycle efficiency than traditional Rankine and gas Brayton power cycles [13]. The combination of performing the cycle in the supercritical state, lower operating pressure ratios, and relatively high turbine exit temperatures allows for extensive recuperation, resulting in a highly efficient and compact power block.

Compact microchannel heat exchangers are commonly utilized in sCO2 power plants, among the different heat exchanger layouts available.The heat exchanger is an effective device to increase the energy efficiency of a process by recovering the wasted energy as much as possible, by transferring heat between two or more fluids for both heating and cooling industrial processes. With the rising effort to make sCO2 cycle more energy efficient, a new generation of heat exchangers is needed that can give greater performance while keeping the heat exchanger's cost and size to minimum. Increasing the surface area to volume ratio of heat exchangers by using mini and micro channels is one way to improve heat transfer performance.130Nowadays, compact heat exchangers (CHEs) are widely used for the heat exchanging process and provide benefits over conventional (shell-and-tube) heat exchangers in compactness, large surface area per volume ratio, and low temperature difference.

CHEs are distinguished by their relatively high area density. The ratio of heat transfer surface to heat exchanger volume is known as area density. According to Shah [14], the compact heat exchanger has a surface area density more than 400 m2/m3 for liquid heat exchangers whereas the surface area density of a typical process industry shell-and-tube heat exchanger is less than 100 m2/m3 [15]. Human lungs are one of the smallest heat exchangers, with an area density of roughly 17,500 m2/m3, or equivalent to about 0.19 mm diameter tubes [16]. The need to reduce energy consumption, capital investment, and component adaptability has sped up research on CHEs and their applications in a variety of fields, particularly in high-temperature services. The high aspect ratios, thin walls, and tight spacing lead to complex geometries that can be difficult to achieve for manufacturers.

The high aspect ratios, thin walls, and tight spacing lead to complex geometries that can be difficult to achieve for manufacturers. As well, the selection of a material compatible with high-temperature applications is also a challenge.Ceramic materials have a number of advantages over typical metallic materials in CHE construction, including high-temperature stability, low material cost, and superior corrosion resistance [16]. The operating temperatures of ceramic based CHEs are significantly higher than those of typical metallic alloys. Ceramic CHEs provide excellent fouling resistance, corrosion resistance, and chemical erosion resistance at high working temperatures. Zirconium diboride is an early transition metal boride that has very high melting points (>3000oC) and is classified as an ultra-high temperature ceramic (UHTC).

Zirconium diboride is an early transition metal boride that has very high melting points (>3000oC) and is classified as an ultra-high temperature ceramic (UHTC). ZrB2 has the lowest theoretical density (6.09 g/cm3) in the family of ultra-high temperature ceramics as well as a unique combination of properties including high strength, hardness, electrical and thermal conductivities, good oxidation, and resistance to chemical attack, which makes ZrB2-based ceramics an attractive candidate for aerospace applications such as leading edges and thermal protection systems for reusable atmospheric re-entry vehicles, hypersonic flight vehicles, and rocket propulsion systems as well as components for corrosive environments and wear-resistant applications [17,18].As a refractory boride, ZrB2 oxidizes when exposed to air at elevated temperatures.

In addition, because of its strong covalent bonding and low self-diffusion coefficient, SiC particles are added to monolithic ZrB2 to increase the densification, mechanical properties, and oxidation resistance. In this study, ZrB2 ceramics containing 10, 20, and 30 vol% SiC particulates were prepared and the four-point bend strength, fracture toughness, elastic modulus, and hardness were measured [17]. The measured data are given in Table 1. Though the elastic modulus and hardness did not change much with the content of SiC, a substantial increase in strength and toughness was noticeable as SiC content increased, with the relative density maintained at ~99%.Table 1: Properties of ZrB2-SiC ceramicsSiC content (vol%) 0 10 20 30Relative Density (%) 99.8 93.2 99.7 99.4Modulus (GPa) 489 450 466 484Hardness (GPa) 23±0.9 24±0.9 24±2.8 24±0.7Strength (MPa) 565±53 713±48 1003±94 1089±152Toughness (MPa.

m0.5) 3.5±0.3 4.1±0.3 4.4±0.2 5.3±0.5131From Table 1, it can be seen that ZrB2 with 30 vol% SiC would be a good candidate for excellent thermal ablation, thermal shock, and oxidation resistance at ultra high temperatures. Densification is needed to prepare the material to achieve the desired mechanical properties. Figure 1 is an SEM image showing the microstructure of ZrB2-30 vol% SiC ceramic material prepared by the Ceramic On-Demand Extrusion (CODE) process.Figure 1: Microstructure of ZrB2- SiC (70/30 in vol%) ceramic material by the CODE processFabricating the compact heat exchanger by additive manufacturing is one of the promising solutions to overcome the manufacturing challenges while considering ceramic as the build material. The method of producing a desired part layer by layer is known as additive manufacturing [20,21,22].

The method of producing a desired part layer by layer is known as additive manufacturing [20,21,22]. The Ceramic On-Demand Extrusion process (CODE), with the constructed machine shown in Figure 2, is a freeform extrusion fabrication process developed at the Missouri University of Science and Technology [19, 20] for producing dense ceramic components. An aqueous paste made of ceramic materials with a very low binder content is used as the printing material in this process. A nozzle attached to the extruder is used to print the desired part geometry layer by layer. The paste is extruded from the nozzle using an auger extruder. The auger extruder uses a syringe in which a material pressurized by compressed air is connected as a barrel. This preloaded pressure is required to deliver the paste to the auger chamber and maintain a continuous flow of the paste. Extrusion is performed by rotating the auger using a servo motor.

This preloaded pressure is required to deliver the paste to the auger chamber and maintain a continuous flow of the paste. Extrusion is performed by rotating the auger using a servo motor. The flow rate is controlled by adjusting the auger’s angular velocity. By maintaining a constant rotation speed of the motor, a constant extrusion rate is achieved.(a)(b)Figure 2: The Ceramic On-Demand Extrusion (CODE) machine132The objective of the current study is to fabricate microchannels with a channel diameter of 500 µm by the CODE process with ZrB2-SiC (70/30 in vol%) material for ultra-high temperature compact heat exchangers. An aqueous paste of ZrB2-SiC is developed with high solid loading that is suitable for printing via the extrusion process. Mechanical properties of the printed samples with this paste are measured, and the printed parts with microchannels are analyzed.2. Experimental Procedure2.1.

Mechanical properties of the printed samples with this paste are measured, and the printed parts with microchannels are analyzed.2. Experimental Procedure2.1. Paste PreparationThe preparation of the paste for extrusion and maintaining its consistency is important since a little amount of fluctuation can change the properties of the paste substantially. The paste for the extrusion contains ZrB2 powder, SiC powder, DARVAN® 821A (Ammonium polyacrylate, Vanderbilt Minerals LLC, Gouverneur, NY), Methocel (Methylcellulose, Dow Chemical Co., Pevely, MO) and deionized (DI) water. The ZrB2 powder has the particle size of 1.5 - 3.0 microns and the SiC powder has an average particle size of 0.65 microns (according to the manufacturer).At first, attrition milling is performed on the ZrB2 and SiC powder particles to mechanically reduce the particle size by milling in coarse milling media with higher rpm.

After that, rotovaping is performed to evaporate all the acetone from the powder that is added in the attrition milling procedure. Batches of ceramic paste are produced by mixing 45 vol% ZrB2-SiC milled powder, DARVAN® 821A as a dispersant agent, and DI water. The mixture is then ball-milled for 24 hours to prepare a homogenous slurry. Methocel is added as a binder in the slurry through a planetary mixer and left for 8 hours to increase the slurry’s density and viscosity into a paste. A syringe charger, THINKY [ARC-40H], is used to load the prepared paste in a syringe which is then attached to the extruder of the CODE machine.2.2 Fabrication of Microchannels for Compact Heat ExchangerComputer-aided design models are made with the desired geometry of the heat exchanger. The primary attempt is to print a heat exchanger with 2x2 channels and to collect data for optimization of paste quality and printing parameters.

The primary attempt is to print a heat exchanger with 2x2 channels and to collect data for optimization of paste quality and printing parameters. Parts with 10x3 and 30x3 microchannels are printed, with the diameter of each channel being 500 µm and the channel length being 50 mm. The surface area density of the 30x3 part is 390 m2/m3. The CAD images are shown in Figure 3.(a)(b)Figure 3: Computer-aided designs of printed specimens: (a) 2x2 channels and (b) 10x3 and 30x3 channels.133The generated CAD file is converted into .stl format, which is then transferred into a slicing software to convert the three-dimensional geometry into multiple layers for the extrusion process to print the specimens layer by layer. The layers then are turned into G-code to create toolpaths for the extruder to follow during printing.Creating microchannels is the most challenging part of printing a compact heat exchanger because of the compact size and good precision needed.

The printing parameters such as printing speed, layer height, and line width need to be optimized in order to achieve microchannels of the desired shape and diameter.(a)(b)(c)Figure 4: Printing microchannels (a) a 10x3 part being printed, (b) a 30x3 part being printed, (c) top view of printed microchannels2.3 Post-Processing of Printed SamplesAfter the parts are printed, post-processing steps including bulk drying, binder burnout, and sintering are needed. Bulk drying is done by putting the printed parts in an environmental chamber with controlled humidity (75%) and temperature (25oC) for 12 hours. By keeping the drying chamber's humidity high and the part’s drying rate low, warpage and crack formation are prevented, resulting in the production of crack-free green parts.

By keeping the drying chamber's humidity high and the part’s drying rate low, warpage and crack formation are prevented, resulting in the production of crack-free green parts. The printed part is easily detached from the substrate after bulk drying.Figure 5: Sintering profile of the printed parts134The binder burnout process is performed in a small retort furnace (Lindberg, St. Watertown, WI) under flowing argon with a heating rate of 5 °C/min to the temperatures of 100, 225, 275, 310, 370, 500, and 600 °C, with a hold of one hour at each temperature before cooling down to room temperature at 10 °C/min. Conventional pressureless sintering in a graphite furnace (Thermal Technology Inc., MODEL#1000-4560-FP30) is done for full densification of the printed ZrB2-SiC parts. As shown in Figure 5, the furnace is run under vacuum at the rate of 20 °C/min to 1450 °C and 1650 °C with a 1-hour hold at each temperature.

As shown in Figure 5, the furnace is run under vacuum at the rate of 20 °C/min to 1450 °C and 1650 °C with a 1-hour hold at each temperature. After the 1-hour hold at 1650 °C, the densification process occurs at flowing argon gas, and the temperature run up to 2100 °C with a 3-hour hold. After this last hold, the samples are cooled in argon to room temperature at the cooling rate of 40 oC/min.2.4 Characterization of Sintered SamplesThe density of sintered samples is measured by the Archimedes method, with water as the submersion medium according to ASTM Standard C373. The dry mass is measured for each specimen, and the samples are then saturated by being immersed in distilled water and kept under vacuum for 12 hours. The bulk density is then calculated by using the saturated and suspended masses.

The bulk density is then calculated by using the saturated and suspended masses. The relative density of the specimen is calculated by dividing the bulk density by the theoretical density.Scanning electron microscopy data are collected by using the TFS ESEM Prisma with Pathfinder EDS, WDS, and EBSD. Prior to the SEM imaging, selection of sectioning, mounting to epoxy, grinding, and polishing of the cross-sections of samples (without etching) are performed. The SEM images are analyzed using ImageJ software (Image J 1.53n, United States National Institute of Health, Bethesda, MD).3. Results and DiscussionThe mechanical properties of the sintered samples are determined, and the quality of the printed microchannels is analyzed.

Results and DiscussionThe mechanical properties of the sintered samples are determined, and the quality of the printed microchannels is analyzed. Figure 6(a) shows a sintered part with 10x3 microchannels and Figure 6(b) shows the cross-section of a part with 20x2 microchannels, where the microchannels are clearly visible.(a)(b)Figure 6: Sintered final parts with microchannels: (a) a sample with 10x3 microchannels, and (b) cross-sectional view of a 20x2 sample135Table 2: Mechanical properties of ZrB2 – 30 vol% SiC samplesFlexural Strength (MPa)Young’s Modulus (GPa)Relative Density (%)ZrB2-SiC (70/30)270 ± 42482 ± 2096Table 2 contains the mechanical properties of the printed parts with ZrB2 – 30 vol% SiC paste. The properties are measured in the printed samples without microchannels. ASTM Standard C1161, with standard type A bars of 2.0 mm x 1.5 mm x 25 mm placed in a screw-driven test frame (model 5881, Instron, Norwood, MA) to determine the properties.

ASTM Standard C1161, with standard type A bars of 2.0 mm x 1.5 mm x 25 mm placed in a screw-driven test frame (model 5881, Instron, Norwood, MA) to determine the properties. The preliminary measurements are conducted with 10 sample bars. The relative density has been measured by the Archimedes method, with water as the submersion medium according to the ASTM Standard C373.The relative density is comparatively lower than the hot-pressed samples shown in Table 1. However, Young’s modulus is similar to that of the hot-pressed samples. The process parameters can be optimized by improving the homogeneity of the paste with higher quality shear mixing procedure during paste making. From a printed green part to a sintered part, the part goes through shrinkage in all three principal dimensions, i.e., along the X-axis (length), Y-axis (width), and Z- axis (height) in Figure 2.

From a printed green part to a sintered part, the part goes through shrinkage in all three principal dimensions, i.e., along the X-axis (length), Y-axis (width), and Z- axis (height) in Figure 2. The shrinkage rates are measured for the sintered parts with microchannels and the average of shrinkage rates in width, length, and height are given in Table 3.Table 3: The average shrinkage rates of the sintered parts with microchannelsSintered ZrB2-SiC sampleWidth 19.01%Linear Shrinkage Rate (%)Length 19.78%Height 22%Volumetric Shrinkage 49.5%SEM micrographs shown in Figure 6 represent the microstructure of some sintered samples with microchannels. Some small pores were observed, which is expected based on the density measurements. The mean cross-sectional area of microchannels calculated is 197,742 µm2, which is essentially the same as the nominal cross-sectional area of 196,350 µm2.

The mean cross-sectional area of microchannels calculated is 197,742 µm2, which is essentially the same as the nominal cross-sectional area of 196,350 µm2. The standard deviation calculated is 27,419 µm2.136Figure 7: SEM micrograph showing the cross-sectional views of various microchannelsFrom Figure 7, it is clearly visible that the cross sections of the microchannels are not circular in shape. This demonstrates that creating circular microchannels by an extrusion-based 3D printing method is a major challenge because of the extremely small channel diameter and also the staircase effect of layered printing. We are confident that significant improvement in creating circular microchannels is possible by using a smaller nozzle diameter. However, associated with the reduction of nozzle diameter could be a large increase in the printing time because cutting the nozzle diameter by one half will require quadruple in the printing time.

However, associated with the reduction of nozzle diameter could be a large increase in the printing time because cutting the nozzle diameter by one half will require quadruple in the printing time. This is a necessary trade- off between part accuracy and printing time. Using a fugitive material to fill in the microchannels and removing the fugitive material in the post-processing may also help making the microchannels more circular. These tests for increasing part shape fidelity will be included in our future study.The continuity inside each channel is also a major concern for heat exchangers. The channel should be entirely clear throughout the whole channel length. The continuity of the microchannels is also tested for the printed parts. Printed samples are cut into multiple parts across the width and light is transferred from one end to the other end of each sample. There is no blockage inside the channels as shown in Figure 8.

Printed samples are cut into multiple parts across the width and light is transferred from one end to the other end of each sample. There is no blockage inside the channels as shown in Figure 8. The pictures may not give the exact geometrical shape of the microchannels cross-section due to possible misalignment during EDM cutting. For parts with larger dimensions, there is a possibility that a few numbers of channels can get blocked due to uneven drying between each layer and surface warping.Figure 8: Testing of continuity in microchannels by using an optical microscope1374. ConclusionsAn aqueous paste of 70 vol% ZrB2 – 30 vol% SiC is developed and the Ceramic On- Demand Extrusion (CODE) process is used to fabricate microchannels for the ultra-high temperature ceramic heat exchanger. The paste is prepared with solid loadings of 45 vol%, which shows a promising outcome for use in the 3D printing process.

The paste is prepared with solid loadings of 45 vol%, which shows a promising outcome for use in the 3D printing process. The relative density of the final (sintered) parts is comparatively lower than the parts prepared by the hot-press method, though Young’s modulus is similar to that of the hot-pressed samples. Microchannels that are 500 µm in diameter are fabricated, followed by the post-processing steps of bulk drying, debinding and sintering. The optical microscope observation of various cross sections cut from the printed samples confirms the continuity inside the printed microchannels. The shrinkage rates of the final parts are measured to be 19-22% in the width, length, and height directions. The cross-sectional shape of the microchannels is not nearly circular. The shape fidelity can be improved significantly by using nozzles of smaller diameters, but with the trade-off of increased printing time.

The shape fidelity can be improved significantly by using nozzles of smaller diameters, but with the trade-off of increased printing time. Using a fugitive material may also possible improve the microchannel fidelity.AcknowledgmentsThe information, data, or work presented herein was funded in part by the Advanced Research Projects Agency-Energy (ARPA-E), U.S. Department of Energy, under Award Number DE-AR0001125. The views and opinions of the authors expressed herein do not necessarily state or reflect those of the United States Government or any agency thereof.References1. Manente, G., & Fortuna, F. M. (2019). Supercritical CO2 power cycles for waste heat recovery: A systematic comparison between traditional and novel layouts with dual expansion. Energy Conversion and Management, 197, 111777. https://doi.org/10.1016/j.enconman.2019.1117772. Kim, Y. M., Sohn, J. L., & Yoon, E. S. (2017). Supercritical co 2 rankine cycles for waste heat recovery from Gas Turbine.

lective Laser Sintering and Fused Deposition Modeling Processes For Functional Ceramic PartsE. Alair Griffin, and Scott McMillinLone Peak Engineering, Inc. 12660 South Fort Street Draper, DT 84020-9362 Tel: 801-553-1732 F~: 801-553-1734 e-lnail: lonepeak@xmission.comIntroductionA variety of solid freeform fabrication (SFF) techniques have been developed toproduce prototype parts directly from a computer-aided drawing (CAD) without any hard tooling, dies or molds [1]. Most of these techniques use polymer, wax, or paper materials to produce the parts.These techniques, with SOlne lnodifications, can be used to rapidly prototypefunctional ceramic parts. Once developed, these techniques could also be used to manufacture small quantities of ceralnic parts on a just-in-time basis.

Once developed, these techniques could also be used to manufacture small quantities of ceralnic parts on a just-in-time basis. Fabrication using conventional techniques is a costly, titne-consuming, and inflexible process when a few ceramic prototypes or when small quantities of parts are needed.Solid freeform fabrication of ceralnic parts offers numerous advantages over conventional processing. Prototypes can be prepared rapidly and cost-effectively. Design changes can be made easily and inexpensively. Larger nUlnber of design options can be investigated. Parts can be designed and engineered to take advantage of the stronger properties of ceramics, while minimizing the weaker ones. Typically, ceramic parts are made using an existing design, regardless of the material used for the original part.

Typically, ceramic parts are made using an existing design, regardless of the material used for the original part. The ability to rapidly prototype a ceralnic component will contribute to concurrent engineering, a popular design process being used today.Lone Peak Engineering (LPE) is developing three SFF processes for ceramics based on:1. Selective Laser Sintering (SLS) 2. Fused Deposition Modeling (FDM) 3. Lalninated Object Manufacturing (LOM™)TMLOM is a registered trademark of Helisys, Inc. Torrance, CA25This paper discusses preliminary results with the SLS and FDM processes. LPE's ceramic LOM based process has been reported at this symposium [2] as well as at other meetings and in different publications [3].ProcedureA three-step procedure was used to fabricate ceralnic components by the SLS and FDM processes: Material Preparation, Green Part(s) Formation, and Heat Treatment.

A flowsheet is shown in Figure 1.Material PreparationThe same procedure was used to prepare the Inaterials for both the SLS and FDMprocesses. The ceramic powdera was mixed with a binder system that contained various binders, plasticizers, and dispersants. The mixing was performed in a sigma blade mixer, which is jacketed, so the mixture can be heated. These mixtures contained between 45 and 55% ceramic powder on a volume basis.Green Part(s) FornlationThe green, unsintered parts were formed using conventional SLS and FDM equipment. The ceralnic-binder mixture was first processed into the proper feed stock for each SFF process. The feed stock was then used to prepare the green parts. Specific procedures used for the SLS and FDM processes are discussed below.SLS ProcessThe cerarnic-binder mixture was ground into a powder using conventional milling techniques. The ground powder consisted of larger particles containing the fine ceramic powder particles in the binder Inatrix.

The ground powder consisted of larger particles containing the fine ceramic powder particles in the binder Inatrix. The average particle size of the ground powder was between 50 and 100 ~lm.The green parts were fabricated in a conventional SLS lnachineb. The ceralnic binder powder was substituted for the nonnal powders used in the SLS lnachine. The green SLS parts are shown in Figure 2. Multiple parts were produced in each run.a A-16SG alumina, ALCOA, Pittsburgh, PA b SLS 125, DTM, Austin, TX26Material Preparation Ceramic-Binder Mixtures+ Green Partes) Formation~ IFilament Preparation I ~ FDMProcess l~ I Powder Preoaration l ~ SLS Process1Heat Treatment Binder Removal SinteringFigure 1. Flowsheet for ceramic-based SLS- and FDM-process.FDM ProcessThe ceramic-binder mixture was extruded into filaments for the FDM process.The filaments were approximately 0.070" in diameter.

Flowsheet for ceramic-based SLS- and FDM-process.FDM ProcessThe ceramic-binder mixture was extruded into filaments for the FDM process.The filaments were approximately 0.070" in diameter. The filaments were used to fabricate the green parts shown in Figure 3 using a conventional FDM Machinec parts in Figure 3 were produced using short discontinuous filaments 6 to 8" long. A single part was produced in each run.. TheHeat TreatmentThe SLS- and FDM-produced parts under went similar heat treatments. The firstheat treatment removed the binder by slowly heating the part in air up to 600°C. After the binder was removed, the parts were sintered in air at 1550°C for two hours.c 3D MODELER, Stratasys, Inc., Minneapolis, MN27Figure 2. Green ceramic parts made by the SLS process.Results and DiscussionBoth the SLS- and FDM-produced parts were sintered without any obviousdegradation or distortion to the parts. The green and sintered properties are shown in Table 1.

The green and sintered properties are shown in Table 1. The results for each process are discussed below.SLSPartsThe sintered density of the SLS parts ranged from 53 to 65 % of theoretical densitl (%TD) with 36 to 47 % open porosity. The green density will have to be increased to achieve a higher sintered density. The green densities shown in Table 1 are between 30 and 35 %TD. The green density will have to be increased to more than 45 %TD to obtain sintered densities above 90 %TD.The variation in green density of the SLS parts may be due to the different shapes of the parts made. Generally, sintered density increases with green density, but this trend was not observed for the SLS parts shown in Table 1. This may also indicate a dependence on part shape.The porous nature of the sintered SLS parts indicates that this process may be an ideal process to produce ceramic cores for investment casting.

This may also indicate a dependence on part shape.The porous nature of the sintered SLS parts indicates that this process may be an ideal process to produce ceramic cores for investment casting. For structural applications the density of the SLS parts will have to be increased.d For this grade of alumina, a theoretical density of3. 90 g/cm3 was used.28Figure 3. Green ceramic parts produced by the FDM process.FDMPartsThe FDM process resulted in parts with sintered densities above 96 %TD. These high densities are a direct result of the high green densities of the FDMed parts. Ceramic components with high densities above 90 %TD such as the FDMed parts could be used in structural appIications.The ceramic-binder systetn used in the FDM process can also be used in an injection molding process. By using the same ceramic-binder system, the FDMed prototype parts will be very similar to the injection molded parts.

By using the same ceramic-binder system, the FDMed prototype parts will be very similar to the injection molded parts. As the volume of parts outgrows the capacity of the FDM process, production can be easily transferred to an injection molding process.Conclusions and ChallengesBased on the prelitninary work performed, the feasibility of the SLS and FDMprocesses to produce ceramic cotnponents has been detnonstrated. The challenge for the SLS process will be to increase the sintered density, if the SLS parts are to be used in structural applications. The challenge with the FDM process for ceramics will be to develop a ceramic-binder mixture that has all of the characteristics necessary to be used in FDM equipment.Acknowledgments29Lone Peak Engineering would like to thank Dr. William Coblenz, from ARPA forsupporting this work under contract # DAAHOl-94-C-R324. LPE also gratefully Table 1.

ramic Structures by Selective Laser Sintering of Microencapsulated, Finely Divided Ceramic MaterialsN. K. Vail and J. W. Barlow Department of Chemical Engineering The University of Texas at AustinAbstractThe feasibility of producing ceramic green parts by Selective Laser Sintering from microencapsulated, finely divide ceramic powders has been reported in an earlier paper. Post-processing of a silica/zirconium orthosilicate system and an alumina system, both utilizing a polymer binder in the form of a coating, are discussed in this paper. Ceramic green parts require post processing to remove the intermediate polymer binder and to impart strength properties to the ceramic bodies. In this paper, the use of ceramic cements and high temperature firing to realize strengths will be discussed.

In this paper, the use of ceramic cements and high temperature firing to realize strengths will be discussed. The effects of cement concentration and controlled drying rates on the strengths and dimensional accuracy of the ceramic bodies are also discussed.Introductionrecent years, in order for manufacturers to remain competitive, the need for the capability to rapidly prototype and develop objects has greatly increased. Concurrently, approaches have emerged to realize rapid prototyping and development [1]. Most are processes which work by the layerwise addition of material. Polymers, with various properties, are the most predominant materials used by these processes although some are attempting to use high temperature materials such as ceramics and metals.

Polymers, with various properties, are the most predominant materials used by these processes although some are attempting to use high temperature materials such as ceramics and metals. While processes using the later materials are a few years away from fully realizing their goals, those processes which use polYmers are making rapid progress in areas of application and resultingquality.Selective Laser Sintering (SLS) is one of these processes which has gained prominence in the area of rapid prototyping within the past few years. This process has its usefulness is not limited to these materials. SLS has been successfully applied to other materials systems which either use an intermediate low-temperature binder or use direct sintering of high-temperature materials [2,3].extremely suitable to the handling of polYmer materials.

However,The use of low-temperature binders, namely, polYmer binders in ceramic systems, has been described and proven effective as a means of producing shapes with good green strength, improved surface finish, and fine feature definition [4]. These "green" parts, however, require post-processing to remove the polYmer binder and to impart greater strength. Greater strengths can be achieved by high-temperature firing to sinter the ceramic material. This method involves a certain amount of densification which results in shrinkage Since most of the ceramic green pieces produced by SLS currently exhibit low of relative densities, on the order of 40-50%, shrinkages involved to achieve desired strengths are large.An alternative to high-temperature sintering is the use of ceramic cements to achieve strength.

Infiltrating a green structure with a ceramic cement then processing it to remove124the polymer binder and cure the cement, is the proposed method of producing ceramic parts with good strengths and little dimensional change.Materials and MethodsA sample of UCAR 430 Acrylic Polymer Latex was obtained from Union Carbide Corporation. UCAR 430 Latex is a styrene-methylmethacryate copolymer paint base. A flow sample of a polymethylmethacrylate (PMMA) characteristics was provided by Rhome-Haas. The properties of these polymers are listed in Table 1. Silica and zircon were obtained from Ransom and Randolf, Inc. as Rancosil #4 and Zircon Flour 325, respectively. A sample of alumina was provided by Lanxide Corporation. The ceramic characteristics are listed in Table 2.latex exhibiting specific meltThree grades of inorganic ceramic cements were obtained from Aremco Products, Inc. These were Cerama-BindTht products grades 542, 643, and 644.

These were Cerama-BindTht products grades 542, 643, and 644. These binders are primarily silicate colloids which, upon curing, provide temperature resistances to 1750°C. In addition to these binders, three samples of Ludoxfl) Colloidal Silica were provided by E.I. duPont de Nemours & Co., Inc. These colloids and their properties are listed in Table 3.Table 1. Polymer properties.Polymer Latex UCAR430 PMMAParticle Size, (nm) 320 290Solids, (wt. %) 45 50Tg' eC)45 93Mw200000 2000Melt Index, (gf10min) N.A. 20.0Table 2. Ceramic powder characteristics.MaterialRancosil #4 Zircon FlourPurity, (%) 99.8 99.8Particle Size, (Hm) 63.0 17.0Density, (gIcc) 2.25 4.56Table 3. Ludoxtll Colloidal Silica propertiesGradeSM TM AMTypeSilica Silica Silica/AluminaParticle Size, (nm) 7 22 12Solids, (wt. %) 30 50 30Three samples of polymer coated ceramic materials were prepared using the method reported previously [4].

%) 30 50 30Three samples of polymer coated ceramic materials were prepared using the method reported previously [4]. These samples included silica/zircon coated with UCAR 430, silica/zircon coated with PMMA, and alumina coated with PMMA. The material make-up and coating conditions for these samples are listed in Table 4. Samples #34 and #35 were admixed with pure ceramic materials. to reduce their respective polymer contents to 15% vol (7.5 wt. % and 5.0 wt. %, respectively). This was done to further reduce the polymer content of the material, to enhance powder packing, and to introduce uncoated particle surfaces for interaction with ceramic cements during penetration.

This was done to further reduce the polymer content of the material, to enhance powder packing, and to introduce uncoated particle surfaces for interaction with ceramic cements during penetration. Admixing of polymer coated and uncoated particles was shown to improve observed bed densities without125significantly affecting the strength of the resultant SLS green part as long as the pure particle content was kept low [4].Sample #283435a e T bl4 CoMaterialSilica Zircon UCAR430 Silica Zircon PMMA Alumina PMMAate ceramIc power rna e-up an coat1n~ con ItlOns.d Content, (vol.) wt. 60.0 62.8 30.0l 14.7) 10.01 22.5) 60.01 65.0) 30.0 (15.2) 10.0 (19.8) 90.0 (80.0 10.0 (20.0k Feed Solids (wt. %)ddTemp. (OC)OutletInlet150605082175501755088Atomizer Speed, rpm375003750037500Selective Laser SinteringAll materials were SLS processed using an SLS1'M Model 125 equipped with a 25 watt CO2 laser.

SLS processing occurred in a nitrogen environment and with elevated bed temperatures. SLS processing conditions for the three powder samples are listed in Table 5. test bars with dimensions 1"x5" and 1''x3'' with total thicknesses ranging from 20-80 layers were produced as were square parts with dimensions ranging from %" to 2" with thicknesses of ~".Several parts were made from each powder sample. Specifically,Sample2834 35Table 5. Selective Laser Sintering operating conditions Beam P o(Wer, SpaCing, (mil) )Layer Thickness, (mil)Scan Speed, {!ps)Temp. eC) ' . AirBed 40650, 70 80125, 62, 5, 81509010 105 3,42 2100 20-10070 35Post-Processing of SLS PartsSLS processed parts were brushed to remove loose powder and their densities were determined. Test samples were subsequently infiltrated with a ceramic cement material. Infiltration was performed simultaneously by capillary action and by dripping the infiltrating material onto part edges.

Test samples were subsequently infiltrated with a ceramic cement material. Infiltration was performed simultaneously by capillary action and by dripping the infiltrating material onto part edges. Once a part had been infiltrated it was allowed to dry for an extended period at ambient conditions. When drying was complete residual moisture was removed by firing for one hour at 200°C. Polymer binder was then removed by firing for an additional one hour at 400°C. Following removal of polYmer binder, parts were re infiltrated with cement material, dried, and residual moisture removed by firing for an additional one hour at 200°C. The mass of each part was measured at the end of each step during the post-processing. Following the second curing, the strengths of several parts were determined by four-point bend analysis. Other samples were fired at high temperature using the firing cycle shown in Figure 1.

Following the second curing, the strengths of several parts were determined by four-point bend analysis. Other samples were fired at high temperature using the firing cycle shown in Figure 1. Strengths of these fired parts were also determined by a four-point bend analysis.12612001000800oi I- ::::l0I-600(j) Q.E 400 (j) f-2000051015 Time. (hours)202530Figure 1. High temperature firing cycle.Results and DiscussionAU test parts made from the three powder samples had sufficient green strengths to permit handling and exhibited good surface finishes as well as good edge definition. Not surprisingly, test parts made from the alumina powder sample exhibited finer features due to its smaller particle size. Test bars made from each of the three powder samples, #28, #34, and #35, were determined to have relative densities of 39.5 ± 1.4, 41.3 ± 1.4, and 45.1 ± 1.2, respectively. Densities increased slightly with increased part thicknesses.

Densities increased slightly with increased part thicknesses. Penetration experiments were conducted on parts made from allthree powder samples. Initially, test pieces made from powder sample #28 (UCAR coated silica/zircon) were penetrated with the Cerama-Bind™ materials. However, during the first penetration tests of these pieces, it became apparent that the viscosities and surface tensions of these In a 1"x3" bar, infiltrated by cements were too great to permit effective penetration. the penetration front would scarcely advance ~" before no further capillary action, movement of the front could be observed.The viscosities and surface tensions were altered by diluting the cement materials with water. None of the three grades, however, was stable to the addition of water although the 643 grade took at least 24 hours to gel. Even when diluted to 50% by volume with water the 643 still did not penetrate test samples effectively.

Even when diluted to 50% by volume with water the 643 still did not penetrate test samples effectively. Subsequently, the binders were diluted with methanol to further reduce their surface tension. Only the 644 grade proved stable to the addition of methanol and penetration tests using this solution proved excellent. Test bars penetrated with a 50% by volume solution of 644 and methanol were wetted easily and completely within a few minutes.test bars were processed as described earlier using the binder methanol solution. Three test bars failed after the polymer burnout stage of the post processing, essentially turning to dust when moved. This weakness at the polymer burnout phase was attributed to the polymer coating which apparently leaves too few exposed ceramic surfaces to which the cement can bind. Extreme care was taken to not damage theEight127remaining test bars during post-processing.

Extreme care was taken to not damage theEight127remaining test bars during post-processing. Also, the initial binder penetration step was altered to try to increase the uptake of cement. Further penetrant was drawn though the bar by placing it on dry towel and dripping more solution onto the surface. The dry towel would draw moisture from the bar thus pulling more material from the surface. This was done on all edges of the bar until no further moisture could be drawn from the bar by dry paper.The results of mass changes of these test bars were all very consistent. A typical sample gained approximately 0.22 gig of green material after the drying period. This amount dropped to 0.21 gig of green material after firing at 200°C, indicating the low moisture content of the dried sample. Following polYmer burnout the mass loss was 0.11 gig of green material, corresponding to a polymer content of 9.6 wt. %. This is in agreement with the expected value of 10 wt. %.

Following polYmer burnout the mass loss was 0.11 gig of green material, corresponding to a polymer content of 9.6 wt. %. This is in agreement with the expected value of 10 wt. %. When infiltrated a second time a bar typically gained an additional 0.15 gig of material after drying and firing at 200°C. This results in a typical mass gain of 41% or 0.29 gig of polymer free material. Initial and final axial measurements are summarized in Table 6.Part # 14 15 16 20 21Table6. Summary of test bar dimensions z, (%) -1.32 -1.29 -1.30 -2.00 -1.66x, (%) -2.34 -2.40 -2.37 -4.27 -3.41mass, (%) 39.3 41.6 42.2 40.4 41.2Final Density, (%) 45.7 47.2 46.4 47.6 46.5(~) -2.33 -2.27 -2.46 -4.35 -3.36Effect of Binder Content on Frac ture Strength Silica/Zircon Test Bars Penetrated with Cerama-Bond 6442001.370".......,.iii S 170 ..c (? c ~ 140 ..... (J) ~ :J -;:; 110 o ~ ..... c '00.. I ..-;f80.,:................1.267 "......., if 1.164 ~ '-" ..c (? c ~ (/) ~ ::J .....

(J) ~ :J -;:; 110 o ~ ..... c '00.. I ..-;f80.,:................1.267 "......., if 1.164 ~ '-" ..c (? c ~ (/) ~ ::J ..... u 0 ~ ..... c '0 0.. I '¢1.0610.9580.8550.7520.6490.5460.443500.34028.028.529.0 Binder Content, ( wi 7.)29.530.0Figure 2. Effects of binder content.Figure 2 shows the results of fracture strengths on the parts listed in Table 6. These data clearly show an increase in the resultant strength with increased cement content. The difference between the two sets of data shown in Figure 2 are attributed to the drying128conditions for the test bars. Those bars showing a higher fracture strength were dried on an open counter top while the remaining bars were dried in a vent hood with the vent operating. While all parts showed some warpage after drying, only those dried in the vent hood exhibited pronounced cracking of the exposed surfaces.

While all parts showed some warpage after drying, only those dried in the vent hood exhibited pronounced cracking of the exposed surfaces. It is presumed these cracks resulted from non-uniform shrinkage of the exposed surfaces due to an increased drying rate caused by the induced air flow of the vent. According to Mujumdar [5], in the drying of clay materials fast drying rates result in a case-hardening. Exposed surfaces shrink about the incompressible core and, as the inner material shrinks, the case will be unable to conform resulting in cracks at weak points. Cracking in these parts occurred primarily along layer boundaries. These large cracks apparently significantly reduced the strength of the bars. In this series of experiments, both part failure following polymer binder removal and surface cracking during the drying phase severely affected the resultant part properties.

In this series of experiments, both part failure following polymer binder removal and surface cracking during the drying phase severely affected the resultant part properties. In an attempt to eliminate these features of the post-processing modifications were made to the powder material and to the drying step. The powder material was altered, as described previously, by the addition of uncoated ceramic powder to the coated powder stock. This introduced exposed particle surfaces for interaction with the penetrating ceramic cement. The drying step was altered by drying infiltrated parts in a humidity controlled environment. With this modification the effective drying rate could be changed with the intent to try to reduce or eliminate cracking.Several square pieces made from the two additional powder samples were infiltrated with a Ludox{lj TM 50% by volume solution of 0.3% wt. Dupanol-ME surfactant in water.

Dupanol-ME surfactant in water. These cubes were dried in an 80% relative humidity environment at a temperature of 72°F for two days. No cracking was observed in these pieces. Following polymer binder removal all parts were significantly more rigid than samples previously observed.Test bars, 1"x3"xY4", made from the alumina powder, were infiltrated with the Ludox TM colloid and fully processed as in previous experiments. A portion of the bars were tested for strength. The remaining portion of bars were fired at high-temperature prior to strength testing. A typical bar had a fracture strength of 1.86 MPa (270 psi) before firing and had a fracture strength of 14.55 MPa (2110 psi) after firing for 17 hours at 1000°C. X ray analysis of the fracture surface fired bars indicated the composition of silica to be homogenous throughout the sample. From mass changes the silica contents of these bars were determined to be about 22% wt.

From mass changes the silica contents of these bars were determined to be about 22% wt. Dimensional shrinkages were approximately 1.0% The increase in silica and the small amount of shrinkage resulted in a in all directions. typical relative density of 61.1%.A number of turbine core samples were made from the alumina powder and post processed in the manner just described. A representative core is shown if Figure 3. This part is approximately two inches square and, as can be seen in the figure, the definition of the part is excellent.ConclusionsThe use of ceramic cements has been shown to be an effective path to the production of ~ll strength ceramic pieces from green objects produced by Selective Laser Sintering. WhIle the results are very encouraging there are some issues which must be resolved.

WhIle the results are very encouraging there are some issues which must be resolved. Some ~f.these issues w~ich are currently under investigation are the effects of drying conditions, fInng cycles, the Improvement of penetration, and the type of ceramic cement best suited to the process.1291.2.3.4.5.Figure 3. Turbine blade core., SolidD.L. Bourell, et.al., "Solid Freeform Fabrication: An Advanced Freeform Fabrication Symposium Proceedings, 1, 1..7 (1990). U. Lakshminarayan, "Selective Laser Sintering of Dissertation, The University of Texas at Austin, 1992. Ni..Sn W. Weiss and D.L. Bourell, "Selective Laser Sintering to Intermetalics", Solid Freeform Fabrication Symposium Proceedings, (1991). N.K. Vail and J.W. Barlow, "Effect of Polymer Coatings as Intermediate Binders on Sintering of Ceramic Particles", Solid Freeform Fabrication Symposium Proceedings, 2, 195..205 (1991). A.S. Mujumdar, Ed., 1987, Dekkar).'O ....

ramic-Polymer Additive Manufacturing System for Ultrasound TransducersMark Cheverton, Prabhjot Singh, L. Scott Smith, Kwok Pong Chan, James A. Brewer, Venkat Venkataramani GE Global Research, Niskayuna NY 12309 cheverto@research.ge.comAbstract A modified multi-layer lithography process is used to additively manufacture high-resolution netshape ceramic structures by photopolymerizing a ceramic-polymer slurry using structured light patterns. This paper will discuss the development of a low-cost, high speed manufacturing process method for the fabrication of piezoelectric ceramic transducer elements in the 1-25 MHz frequency range. The key considerations in the development of the process including the selection and optimization of slurry materials for the deposition of PZT 5H materials, optical exposure parameters, debinding/sintering profiles, process monitoring, and post manufacturing processing.

Ongoing work includes improvements to the materials properties, improved throughput and geometric fidelity.IntroductionUltrasound transducers constitute a key sensor technology for both medical and non-destructivetesting. Ultrasound has become the most common medical imaging procedure because the process is fast, avoids ionizing radiation, is relatively portable, and uses inexpensive equipment. For the non- destructive evaluation of high value metal, ceramic, and composite parts, the use of ionizing radiation is less of a concern, while the other characteristics still apply. In both of these applications, the ability to efficiently generate high quality ultrasound beams in the inspected body is critical. For medical imaging that means the use of phased array technology where individual transducer elements, and time delays are used to form the transmitted and received beams. This same technology is used in some non- destructive testing systems.

This same technology is used in some non- destructive testing systems. However, there is also a significant use of single element ultrasound transducers where the use of 1-3 composite transducer structures1 improves beam quality. So in both applications, arrays or composites, the fabrication of small pillar-like structures is central to the device technology. Conventional technology, based on dicing plates of piezoelectric and other acoustic impedance matching and backing materials, has been functionally effective. Because it is based on a serial and sequential process, it does not scale well to higher production volumes and tends to be slow, or equivalently requires parallel capital equipment investment. Additive manufacturing provides a novel alternative approach to fabricating ultrasound transducers.Prior Work: Extensive prior work on the layered deposition of various structural and electrical ceramic materials has been reported in literature.

Several modes of layer creation have been researched. These include863extrusion of semi-molten thermoplastic-ceramic filament material2-3, photocuring of ceramic-polymer slurries4-8 inkjet of wax-ceramic9 mixtures etc. Additionally, micro-molding techniques have also been researched10. [11] summarizes various micro-layered manufacturing processes for ceramic fabrication. In the context of photocurable ceramic-polymer layered deposition processes, key process considerations include the ability to:Formulate low-viscosity slurry materials that can be spread to form 10-100 micron thick layers with high consistency. This is usually accomplished by suspending fine ceramic particles with stearic dispersants in a photocurable (acrylate, epoxy etc) monomer Photocure the slurry layers and form cross-linked polymer chains within and across layers using a high intensity, UV-rich light pattern. This pattern could be formed by a laser or a digital spatial light modulator.

This pattern could be formed by a laser or a digital spatial light modulator. The refractive index of the slurry constituents, the incident light spectrum and the particle size distribution usually affect the depth of cure and the resolution of the patterns Sinter the ceramic-polymer green parts to high density using carefully controlled thermal cycles in the appropriate environmental conditionsRelatively limited work has been reported on the photocuring of lead-based piezoceramic slurries12-13. This is largely due to two factors; first, the high refractive index of such ceramic materials, which causes light to scatter preventing cross-linking through the thickness of the slurry layers; second, the sintering of lead-rich piezoceramics is difficult due to the volatility of lead at high temperatures leading to poor electromechanical properties in sintered structures. The relatively high surface area of netshape green parts amplifies this problem.

The relatively high surface area of netshape green parts amplifies this problem. In the following section we discuss the experimental methods used to netshape manufacture of lead zirconate titanate (PZT) parts in clinically relevant geometries (size and feature fidelity).The paper is organized as follows: system architecture with an emphasis on components and challenges encountered & the material system will be described. Thereafter, the manufacturing process will be discussed, and finally the post processing with lessons learned.Manufacturing Process:The manufacturing process can be summarized with four steps, the dispensing of the photo curable material, the spreading of the material, the exposure of the material, and finally the removal of all uncured material.

This process is summarized in figure 1 below:Figure 1 – Additive process for manufacture of ultrasound transducers.System architecture864A key feature of ultrasound transducers is their constant cross sectional structure, which make alayerwise additive process using a fixed mask for patterning very attractive. The system built at GE’s Global Research Center focused on this layerwise process, using UV curable materials. The system uses high resolution x/y/z stages, automated dispenser to deposit a controlled amount of material, a high intensity UV light source, and a chrome-on-glass mask with the pattern to be written, and multiple cameras to monitor the manufacturing process and detect problems if they occur during build. The entire system was integrated with custom software developed in LabVIEW.

The entire system was integrated with custom software developed in LabVIEW. The following system categories will be discussed in turn: material system, light source, material deposition, material spreading, motion control and the software system, process monitoring, and part post processing.Material system:Starting materials & Slurry PreparationThe current UV-curable PZT formulation was prepared using PZT powder as the ceramic fillerand a diacrylate photocurable resin binder with an appropriate concentration of a photoinitiator. A stearic dispersant was used to promote the dispersion of PZT particle within the organic resin.The ceramic slurry formulation with 80wt% of ceramic content was prepared by mixing the constituents in a planetary mixer for 2 hrs using grinding media to break agglomerates in the ceramic powder. Isopropanol was added to the mixture to adjust the viscosity of the slurry.

Isopropanol was added to the mixture to adjust the viscosity of the slurry. Thereafter the slurry was transferred to a 30cc EFD syringe cartridge for use with the deposition system.A significant effort went into developing a slurry with specific flow characteristics of viscosity ~ 10 poise at shear rate between10 s-1 to 100 s-1 (typical shear rate range of our equipment). This involved mixing various components with zirconia media balls in a high speed planetary mixer to reduce the particle size. A viscosity profile of PZT slurry was shown in figure 2. The slurry exhibits a shear thinning behavior of a non-Newtonian fluid.

A viscosity profile of PZT slurry was shown in figure 2. The slurry exhibits a shear thinning behavior of a non-Newtonian fluid. If the material was mixed and allowed to sit for longer than half a day, then the material tended to not flow or spread very well, and would not make acceptable parts.8651000100) e s i o P (y t i s o c s i V1010.1110 Shear rate (S-1)1001000Figure 2: Viscosity of the photosensitive PZT-diacrylate slurryThe amount of PZT and polymer (HDODA) was determined mathematically, such that the volume fraction of PZT would be 38%-40% after IPA evaporated. Photoinitiator amounts were determined empirically by testing fixed exposure times with varying amounts of photo initiator and assessing the level of adhesion between successive layers. With too little photo initiator, the parts do not cure sufficiently and fall apart when cleaning the uncured slurry out of the green part.

With too little photo initiator, the parts do not cure sufficiently and fall apart when cleaning the uncured slurry out of the green part. With too much photoinitiator, the scattering increased, decreasing pattern resolution with small features becoming merged together.Light source:Knowing that there would need to be high solids loadings in order to achieve the densities necessary for ultrasound transducer applications, the first consideration in selecting the light source was high power. LED fiber light sources were considered, but with the desire to operate in the UV range, there appeared to be limited options to achieve high power and relatively collimated light output. As a result, the Newport model 92292 UV solar simulator was selected for use in this system. This light source yielded a total output of 150 mW/cm2 in the UV-A range and 75 mW/cm2 in the UV-V range.

As a result, the Newport model 92292 UV solar simulator was selected for use in this system. This light source yielded a total output of 150 mW/cm2 in the UV-A range and 75 mW/cm2 in the UV-V range. Communication to this light source was done via RS232.Material deposition:Material deposition was performed using a pneumatic pump, Nordson EFD Ultima model as wellas with Aerotech translation stages, integrated using LabVIEW. This system used air pressure to dispense material from a 22 gauge syringe tip attached to a 30cc tube. To activate the pump, a TTL signal was used, the high state turning on the pump and allowing it to dispense, and the low state turned off the pump.866The slurry was dispensed as a bead across the substrate, with the location of the bead determined experimentally so that the material would be positioned properly. The system would move to the start of the bead, and then the pump turned on.

The system would move to the start of the bead, and then the pump turned on. The motion control system them moved the substrate under the dispenser so that a bead was deposited parallel to the doctor blade. When the system reached the end point, the dispenser was turned off. The volume of material being deposited was controlled by controlling the speed of the stage as it traversed under the dispenser. A slower speed allowed for more material to be deposited. A camera system was designed & installed to detect flow stoppages in the syringe. It was observed that stoppages in the build led to subsequent delaminations in the part.Material spreading:The dispensed bead, as described above, was spread usinga stainless steel doctor blade. This doctor blade was carefully leveled to the substrate using micrometers so that the gap on both ends of the doctor blade were the same.

This doctor blade was carefully leveled to the substrate using micrometers so that the gap on both ends of the doctor blade were the same. The vacuum chuck which held the glass substrate, upon which everything was built, had to also be leveled to the blade, so that the gap between the blade and the substrate would be the same at the beginning and the end of the wipe. This leveling was done with thin shim material so that the gap at the beginning and the end of the wipe would be within 20 um or better.It was found to be important to level the vacuum chuckFigure 3 – surface finish with a wipe speed of 0.5 mm/swith the vacuum on, as the glass substrate changed shape slightly, likely due to the non-flattness of the vacuum chuck.

In any event, if the leveling was done with the vacuum on, and then parts were built with the vacuum off, or vice versa, there was a chance of the doctor blade colliding with the glass substrate, causing the process to fault and stop.One of the challenges that was discovered when conducting experiments on the spreading of the material was the relationship between surface finish and wipe speed. Initially, the wipe speed was set to a small number, in the neighborhood of 0.5 mm/s. This was done so that the wiping would not destroy any of the subsurface features that had already been printed. However, it was found that the slow speed yielded a poor surface finish, as shown in figure 3. In comparison, figure 4 shows the surface finish with a wipe speed of 3 mm/s. These images were acquired by looking at the specular reflection of a structured light. In figure 3, the dark stripes cannot be seen, but on figure 4, the dark stripes are clearly visible.

These images were acquired by looking at the specular reflection of a structured light. In figure 3, the dark stripes cannot be seen, but on figure 4, the dark stripes are clearly visible. Wipe speeds greater than 3 mm/s were tried, but speeds greater than 3 mm/s were found to be detrimental to already printed subsurface structures. To document theFigure 4 – surface finish with a wipe speed of 3 mm/s867quality of each layer, a camera was installed in the system to take a picture of every layer before exposure. This allowed analysis of the quality of the wipe to detect problems like not dispensing enough material, or viscosity changes causing insufficient volumes of slurry being dispensed, which in turn caused incomplete layer coverage. More information about the surface imaging camera will be discussed in the process monitoring section.Custom control softwareThe control of these stages was done using LabVIEW.

More information about the surface imaging camera will be discussed in the process monitoring section.Custom control softwareThe control of these stages was done using LabVIEW. The software was developed with experimentation in mind, but it also had a manufacturing setting, so that control of the system could be limited and operation could be done by individuals with only basic training on how to operate the system. The software was written with a standard-state machine architecture, with an event structure at the beginning of the state machine for experimentation. The series of steps in the software were the following: dispense, wipe, clean, evaporate solvent, and expose.Dispense - In the dispense step, the system moved to the start of the bead that would be deposited. This location was stored in an ini file that could be easily edited with a simple text editor, or it could be modified with tools built into the software.

This location was stored in an ini file that could be easily edited with a simple text editor, or it could be modified with tools built into the software. The start point, the end point, and the velocity of the traverse were defined in the ini file, with lower velocities allowing the system to deposit more material and high velocities for less material.

The speed of the translation stage, the air pressure used on the pump, and the needle size were all factors that were optimized to get the right amount of material deposited so that there was complete coverage without too much excess slurry left over at the end of the wipe.A camera, LED backlight, and mirror were positioned so that the dispense could be monitored.The LED backlight and camera were positioned on opposite sides of the dispense stream, with the mirror used to fold the path so that it fit within the footprint of the system.Wipe - After dispensing the bead, the stages moved so that the doctor blade was just behind thebead of slurry. Typical layer thicknesses for this process were 20 um. As a result, the doctor blade had to be very close to the substrate or previous layers.

Typical layer thicknesses for this process were 20 um. As a result, the doctor blade had to be very close to the substrate or previous layers. To mitigate the risk of the doctor blade overshooting its position and colliding with the substrate, the system moved to within 1 mm of the target height, and then very slowly approached the final position at approximately 0.1 mm/s. This slow approach eliminated any collisions between the doctor blade and the previously printed layers.The start point of the wipe, the end point of the wipe, and the wipe speed were defined in thebuild parameters, in an ini file. The engineer running the system could modify these parameters with a text editor, or use on-screen tools to modify these settings.A critical concern with the wipe was found to be the wipe speed, not just because of surface quality, but also with respect to current draw on the stages.

If a very low wipe speed was used, the z- stage would draw excessive current in order to keep its height constant, frequently causing868unrecoverable over-current faults when the average current draw went over some threshold set in the Aerotech parameter file, causing the process to stop. It was found that the higher wipe speeds not only produced a better surface finish, but also reduced the number of over-current faults with the stages.Clean – After spreading the material over the substrate, the system moved to a location such that a sponge was positioned at one end of the doctor blade. The sponge was then wiped across the doctor blade, to clean any residual slurry left on the blade. If this step was skipped, then one ran the risk of leaving slurry on the blade, and having the slurry cured by collateral exposure from the UV light, destroying the nice, flat surface of the doctor blade.

As a result, this cleaning step was performed immediately after each wipe.Evaporate solvent – Before the exposure, excessive solvent was removed from the part bymoving the part in front of a fan for one minute. During that minute, the fan would be turned on, via LabVIEW, and the residual solvent that was added to the slurry, to lower the viscosity so that it could be dispensed easily, would evaporate, allowing a small amount of shrinkage. During this step, an LED line light would be turned on, and the light would be emitted with a grazing incidence angle, parallel to the surface of the material. A camera would capture an image from above; the glancing lighting allowing topological features to be highlighted and captured by the camera for later analysis. Images from this camera can be seen in figure 7 and 8.Expose – After the evaporation step, the stages would move the sample under the light source for exposure.

Images from this camera can be seen in figure 7 and 8.Expose – After the evaporation step, the stages would move the sample under the light source for exposure. Typically, twenty base layers, without any pattern, would be exposed into the material. After those base layers were complete, then the desired pattern would be exposed by positioning a chrome-on-glass mask from Advanced Reproductions, Inc. between the sample and the light source. This mask was held on a translation stage from Velmex, and moved into position after the base layers were complete.It was important to position the uncured layer as close to the mask as possible, in order to get sharp shadows and good size definition. The spacing chosen for this process was 0.25 mm. As a result, the parallelism between the mask and the substrate was very important.

The spacing chosen for this process was 0.25 mm. As a result, the parallelism between the mask and the substrate was very important. A great deal of effort was put into designing a fixture that would hold the mask but allow enough degrees of freedom to easily get it level to the substrate while still being strong enough so as it would not droop over time. A two axis kinematic mount with micrometers was selected from Thorlabs. The micrometers made leveling much easier than using shims.With such a small gap between the mask and the uncured material, collisions were problematic. Therefore, as described above, the stage would move the substrate to within 1 mm of the final exposure position, and then approach the desired exposure height with a very low velocity. If a collision did occur, then uncured slurry would get on the mask, and then be cured to the mask by the UV light.

If a collision did occur, then uncured slurry would get on the mask, and then be cured to the mask by the UV light. When this would happen, it was necessary to remove the mask completely, soak it in acetone overnight, then glue it back to its frame again, remount it to the Velmex stage, and relevel it to the substrate.869Once the exposure step was complete, the system would then repeat the process, moving thestage to the start of the dispense step, and go through all the steps again. The wipe and the expose steps had to track the height of the part, and change their z-position to account for all previous layers, so that the height of the uncured slurry was always held constant, even though the z-height of the stage was not.To enable experimentation, all of the above steps could be run individually using preset parameters.

As a result, when new materials were being developed, it was common for the researcher to do a test dispense, just to see how the material would flow, and then perhaps do a test wipe, just to see the flow dynamics and surface finish. The screen for the control software is shown on figure 5.Figure 5 – screen shot of control softwareProcess monitoring:Two cameras were integrated into the system, one to monitor the dispense, and a second to monitor the surface quality of the layers just before the exposure. The goal of these cameras was to notify the operator about a problem early in the process, so that corrective actions could be taken to save the build, or terminate the build and start a new one.

The goal of these cameras was to notify the operator about a problem early in the process, so that corrective actions could be taken to save the build, or terminate the build and start a new one. Because these builds could sometimes take many hours to complete, it was important to have the capability to monitor the process and have automated decisioning built into the software system.870Dispense Camera:This camera looked at the dispense stream and verified that something was dispensed. This was done by putting an LED backlight on one side of the dispense stream, and the camera on the other, with a mirror on the camera side to fold the optical path, making the camera fit within the footprint of the system. The system would take a picture of the dispenser before it was actuated, then take a picture during dispensing, and subtracting the two images so that only the slurry stream could be extracted from the image.

Because the slurry is opaque, it appeared black in the image, with the background (the LED backlight) being saturated at a grey value of 255. With such a large contrast, it was easy to detect the stream and verify that the system was actually dispensing slurry.Nozzle clogs were detected and the system would attempt to re-start multiple times. Once the needle on the dispenser was changed and a successful dispense was done, the system would then send an email to the team notifying them that the process had been successful restarted and the problem solved. If the needle was replaced within twenty to thirty minutes and the process continued, then the build would still be successful. Excessive solvent vaporization and water uptake were suspected to be the two leading causes of dispense failures. Figure 6 shows images captured from the dispense camera.

Excessive solvent vaporization and water uptake were suspected to be the two leading causes of dispense failures. Figure 6 shows images captured from the dispense camera. The thicker section at the top of figure 6 (a) is the actual nozzle, and the thick black line at the bottom off the image is part of the drip cup.Surface camera:A megapixel camera was positioned over the sample with a working distance of approximately 18 inches so that the camera was out of the way of the operator during loading and unloading of the system. An LED line light from CCS-America was positioned so that the light from the line light was focused across the top layer of the sample, with the light collimated parallel to the surface. This grazing lighting position was used to highlight one side of a bump on the surface of the part being build, and cause a shadow on the back side of the bump.

This grazing lighting position was used to highlight one side of a bump on the surface of the part being build, and cause a shadow on the back side of the bump. In this way, the lighting highlighted topological features in the slurry layer.Figure 6: a – image of the slurry stream during deposition, b – detected stream after image processingJust before exposure, the motion stages were used to place the deposited layers underneath the camera, defined in the build parameters. LED lights were turned on and an image of the layer was taken and saved for analysis post-build. It was possible to detect topological anomalies and make decisions, like re-wipe the layer, or put on more slurry and wipe again, or even abort the build, however these features were tested but never implemented into the final software version.

Figure 7 shows an image of the sample surface just after the wipe but before the exposure on layer 12.Figure 7 – image from the surface camera.871There are few features to be seen in this image, however, as this build continued, the slurry’s viscosity increased, making the dispense put down too little slurry, creating a topological feature at the end of the wipe. These features can be seen in figure 8.This surface camera proved essential in understanding the kinetics of the wiping process, and the effects that the sub- surface structures can cause on the fluid dynamics of the wiping process.Figure 8 – image from the surface camera.Post processing:Cleaning:The most challenging part of the process was thecleaning of the uncured slurry from the part, without damaging any of the cured features. For the ultrasound transducers being manufactured on this additive system, a series of bars were printed, as shown in figure 9.

For the ultrasound transducers being manufactured on this additive system, a series of bars were printed, as shown in figure 9. The typical dimension of the bars were 250 um wide with a 60 um kerf in the green state. The image in figure 9 is a post sintered structure.Figure 9 – micrograph of ultrasound transducer ceramic green parts.Upon completion of the build process, the build plate (glass) was removed from the vacuum chuck. Excess uncured slurry was removed with a spatula, then the part, still on the glass substrate, is put in an ultrasonic bath to remove excessive slurry. Delaminations and broken features were observed during the cleaning process. Process yield was ~40%. Parts were then dried in an oven.Sintering:A 2-stage densification process was developed to obtain high-density ceramics.

Process yield was ~40%. Parts were then dried in an oven.Sintering:A 2-stage densification process was developed to obtain high-density ceramics. In the first stage,the organic constitutents of the part were pyrolyzed in an oxygen environment using a carefully designed thermal cycle based on the thermogravimetric analysis of the photocurable binder and the surfactant. This is essential to avoid thermally shock which leads to part breakage. Temperature in the first stage was restricted to 700C to avoid lead loss. In the second stage, the part was placed on platinum foil and the sealed in a magnesia crucible with extra piezoceramic source powder. Magnesia sintering furniture was selected to minimize reactions with lead which is volatile above 700C. The extra source powder was added to ensure a positive vapor pressure of lead during sintering. A Zircar Zirconia box furnace was used to sinter the part.

The extra source powder was added to ensure a positive vapor pressure of lead during sintering. A Zircar Zirconia box furnace was used to sinter the part. A uniform heating rate of 1 degree/minute was used to prevent thermal shock and obtain isotropic shrinkage. Density was measured using the Archimedes principle in a water bath of known temperature. Special setters were created to ensure that the features in the ultrasound transducer ceramic stayed straight (Figures 10 & 11).872Figure 10 – transducer after sinteringFigure 12 – sintered transducer without setterTransducer performance:Parts produced by this process are in the 95%- 98% dense, with a grain size of 2-3 microns. Lead loss during sintering was less than 0.5%. This was determined by measuring the weight of the parts before and after sintering. The electromechanical coupling coefficient (kt) was between 0.52-0.53 with a dielectric constant of 2000.

This was determined by measuring the weight of the parts before and after sintering. The electromechanical coupling coefficient (kt) was between 0.52-0.53 with a dielectric constant of 2000. The resulting probe operates at a center frequency of 2.x MHz.The piezoceramic composite was processed into a convex transducer.Figure 11 – transducer after sintering improvement w/ settersFigure 13 – sintered transducer with setterFigure 14 – ultrasound images of a fetal phantom. The image on the left is from the transducer build by the additive process described, and the image on the right was from the commercial ultrasound system.873The ceramic wafers were built into a convex transducer array and tested on a fetal a phantom. The images from the GE constructed ultrasound transducer were compared to images captures from a commercially available ultrasound system.

The images from the GE constructed ultrasound transducer were compared to images captures from a commercially available ultrasound system. The images from these tests can be seen in figure 14.ConclusionThrough the use of an additive process, it was demonstrated that ultrasound transducers couldbe manufactured with high resolution features that yield images comparable to commercial systems. Slurry chemistries and cleaning processes were developed to optimize the yield of the manufacturing process. In addition, machine vision based process monitoring was integrated into the custom system to monitor the process, so that when problems occurred, operators could intervene quickly rather than waiting for the long build to finish, only to yield a failed part.

Future work on this process will include further optimization of the slurry chemistry & improvements to the deposition set-up .This work was supported by Award Number RC2EB011439 from the National Institute of Biomedical Imaging And Bioengineering. The content is solely the responsibility of the authors and does not necessarily represent the official views of NIBIB or the NIH.References1. W.A. Smith The Role of Piezocomposites in Ultrasonic Transducers. Proceedings of the 1989 Ultrasonics Symposium 755-766.2. Safari, M. Allhverdi, E.K. Akdogan, Solid freeform fabrication of piezoelectric sensors and actuators, J. of Materials Science, Vol. 41, pp. 177-198, 2006.3. Safari, M. Ebrahimi, S. Turcu, A. Hall, R. Brennan, N. Hagh, Layered manufacturing for prototyping of novel transducers, In proceedings of Ultrasonics Symposium, 2002.4. X. Zhang, X. N. Jiang, C.

RAMICS 3D PRINTING BY SELECTIVE INHIBITION SINTERINGBehrokh Khoshnevis*, Jing Zhang*, Miranda Fateri**, Zichen Xiao***Department of Industrial and Systems Engineering, University of Southern California, Los Angeles, CA 90089**Department of Mechanical Engineering and Mechatronics, FH Aachen, DE***Department of Aerospace and Mechanical Engineering, University of Southern California, Los Angeles, CAAbstractSelective Inhibition Sintering (SIS) has been proven effective in producing polymeric and metallic parts. Due to the low cost and high quality of SIS printing, the impact of SIS printing in the 3D printing industry could be disruptive. The potential of SIS is further extended to ceramics, an important but hard to print material, by the same mechanism of creating an easy-to-break sacrificial mold. Due to the high sintering temperature of ceramics, fluid based inhibitors delivered by inkjet printing tend to not be effective in SIS for ceramics.

Due to the high sintering temperature of ceramics, fluid based inhibitors delivered by inkjet printing tend to not be effective in SIS for ceramics. Accordingly, the new concept of inhibition by dry powder delivery is implemented. Preliminary experiments have shown the feasibility and ease of printing of simple ceramic parts. Additional experiments are underway to increase the possible part complexity and accuracy, and to optimize the sintering process.Keywords: Additive Manufacturing (AM), Selective Inhibition Sintering (SIS), Sintering Inhibitor, Ceramics 3D PrintingIntroductionAs a new approach for powder based Additive Manufacturing (AM), SIS has been proven efficient in producing 3D printed parts of high quality with polymers and metals [1,2](Figure 1).Figure 1 SIS Printed Metallic PartsCurrent AM methods for metals and ceramics use high density energy (like EBM and SLS) or post sintering of binder-consolidated green parts.

The print time in the current methods depends on the size of the parts hence printing large parts requires significantly longer print time. Furthermore, for printing large parts, the high density energy beam of laser and electron beam lose focus at peripheral regions of layers causing reduced energy density and poor quality.The SIS process is fast and guarantees uniform part quality. As the powder layer is paved (Figure 2), the inhibitor is deposited along the part layer profile, often a 2D border line that shortens the printing time. The inhibitor is deposited layer by layer As the printing process is completed, the green part is moved into a sintering furnace where the inhibitor coating serves as a temporary sacrificial mold, separating the parts from the redundant powder. As the sintering process is completed, the part is taken out from the furnace and the redundant materials are removed by means of sand blasting.

As the sintering process is completed, the part is taken out from the furnace and the redundant materials are removed by means of sand blasting. The fabrication resolution depends upon the precision of deposition of inhibitor and base powder particle size and does not degrade as part size increases.Figure 2 SIS process 1. Powder pavement; 2. Inhibitor deposition; 3. Sintering in the furnace; 4. Part separationThe original SIS technology has been satisfactory for fabrication of polymeric and metallic parts, however, in production of parts made of metals with high sintering temperature as well as in production of ceramic parts (which often require high temperature to sinter), the inkjet-printedliquid inhibition approach seems to have some limitations with respect to ease of separation of the adjacent uninhibited powder regions.

Also for 3D printing in space and on the Moon and asteroids where the vacuum environment evaporates fluidic inhibitor solutions the use of non- fluidic inhibitors is necessary. In fact the need for printing inter-locking ceramic tiles for Lunar landing pads in a NASA supported project has motivated the development of the new SIS process.For additive manufacturing of ceramics, SIS has some obvious advantages over competing technologies. Electron beam cannot be used for ceramics as most ceramics are not conductive. Laser beam works in heating up the ceramic powder but the energy requirement is high.

Electron beam cannot be used for ceramics as most ceramics are not conductive. Laser beam works in heating up the ceramic powder but the energy requirement is high. Finally, binder based sintering methods could lead to high shrinkage rate [3] and the remnants of the binder in the matrix weaken the ceramic parts.Application of SIS upon ceramic parts manufacturing is proven feasible and large volume manufacturing of low cost ceramic parts are possible.Principle of operation of SIS_CeramicsIn the new SIS process the inhibitor material is fine dry powder form of a ceramic (such as magnesium oxide which has a sintering temperature of over 1500 C) or other material with sintering temperature that is higher than that of the base powder material. Controlled delivery of the inhibitor powder within the base material powder of which the 3D part is to be fabricated can be done using a vibrating nozzle.

Controlled delivery of the inhibitor powder within the base material powder of which the 3D part is to be fabricated can be done using a vibrating nozzle. High-resolution deposition of thin walls of inhibitor powder within loose base powder is quite possible with a fine nozzle. Using this approach the desired geometry of the part layer boundary may be created within the base material powder. Note that a rotary axis actuator can be used to pivot the nozzle so that its orifice is always positioned on the opposite side of the nozzle motion direction. The entire nozzle assembly may be moved in the three dimensional space by means of a motion control platform. This way the inhibitor powder may be deposited according to any straight or curved path.Inhibitor material may be deposited for each successive powder layer spread by a powder spreading mechanism until all layers are completed.

The entire loose powder layer vat may then be placed in a powder sintering furnace which may be operated by means of resistive heating, microwave or other heating means that sinters the base powder but leaves the inhibitor powder intact because of its higher temperature sintering heat requirement. After sintering the final part may be separated from the rest of the sintered material. In case of complex part geometries separation lines may be created by the nozzle outside the part geometry connecting selected points on the layer boundary to the edge of the powder volume to ease part separation. A prototype SIS-Ceramic machine is shown in Figure 3.Figure 3 Prototype machine of SIS-CeramicsTo separate the sintered parts from the redundant material, the inhibitor powder must remain loose after sintering. Therefore, the inhibitor needs to have a higher sintering temperature than the ceramic powder that makes the part.

Therefore, the inhibitor needs to have a higher sintering temperature than the ceramic powder that makes the part. The inhibition mechanism (Figure 4) illustrates the inhibition mechanism. As the inhibitor is deposited into the base layer, a thin wall is formed between the inside and outside sections of the base material that constitute the part. Before sintering, both base material and inhibitor material are loose powder. During sintering, the building material gets sintered and hence shrinks and forms the solid part while the furnace temperature is well below the sintering temperature for the inhibitor powder. After the part is taken out from the furnace, the loose inhibitor is easily removed and the desired part is extracted.In our preliminary experiments, lunar regolith simulant (JSC-1A1) is used as the building material and aluminum oxide is used as the inhibitor. The lunar regolith simulant is sifted through a 60 meshes sieve and the powder size is below 250 um.

The lunar regolith simulant is sifted through a 60 meshes sieve and the powder size is below 250 um. The sintering of JSC-1A is carried out at 1150 °C in the ambient environment for 1 hour and then the part is cooled down to the room temperature. The sintering temperature for JSC-1A is about 1100 °C and the sintering temperature for aluminum oxide is about 1400 °C [4]. Near-spherical aluminum oxide powder is used for ease of deposition through the thin nozzle channel.1 http://www.orbitec.com/store/simulant.html(a)(b)Figure 4 Inhibition Mechanism for SIS-Ceramic (a) Simplified powder distribution before sintering; (b) Shrinkage of building material with inhibitor almost untouched. (The black spheres represent the building material and the white spheres represent the inhibitors)Experimental ResultsResults show easiness in separating the printed parts from the redundant material. As can be seen in Figure 5, the block taken out from the furnace has been well sintered.

As can be seen in Figure 5, the block taken out from the furnace has been well sintered. While the area deposited with inhibitor (the white curves in the figure) is not sintered at all. The redundant material sintered can be broken away by hand easily along the separation lines. The inhibitor powder on part surface could brushed away with ease.Figure 5 Sintered Ceramic Parts and Separation from the redundant materialDue to oxidation, the raw black powder has turned red.Analysis Our tests show that the printed parts are strong and have high impact strength. The extreme ease of separation guarantees the feasibility of the process for complex part geometry. Much thinner inhibitor walls may be built to increase part resolution.

The extreme ease of separation guarantees the feasibility of the process for complex part geometry. Much thinner inhibitor walls may be built to increase part resolution. As long as the inhibitor has a distinctly higher sintering temperature than the base material for making the part, the inhibitor would be a qualified candidate for use.ApplicationsThe SIS-Ceramics process may be used to manufacture complex ceramic parts with high quality. The temporary sacrificial mold protects the part from deformation and the built parts can have high strength with no internal impurities. Ceramic parts generally have higher service temperature and are oxidation resistant. SIS printed ceramic parts could resist very high temperature as the process allows the sintering of ceramics with very high melting point. For outer space fabrication of objects such as landing pad tiles the materials need to sustain high temperature exhaust plume of the landers.

For outer space fabrication of objects such as landing pad tiles the materials need to sustain high temperature exhaust plume of the landers. The sintered lunar regolith could well satisfy this purpose. The SIS process described here can utilize in-situ resources to create a variety of ceramic objects in outer space. In Figure 6, interlocking tiles built with lunar regolith simulant are shown.Figure 6 Printed interlocking brick (left) Interlocking pattern for lunar landing pad; (middle)Sintered brick; (right) Printed bricksConclusionSIS-Ceramic has been proven effective and efficient in producing ceramic parts. The fabricated parts are free from contaminants and can be used in a broader fields of application. Our current focus on the improvement of resolution using novel mechanisms for inhibitor powder delivery.References[1] B Khoshnevis, M Yoozbashizadeh, Y Chen. Metallic part fabrication using selective inhibition sintering (SIS), Rapid Prototyping Journal.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCertification Strategy for Additively Manufactured Structural FittingsRobert M. Taylor 1 University of Texas at Arlington, Arlington, TX, 76019Joe Manzo 2 Titan Industries, Tempe, AZ 85281Lori Flansburg 3 Lockheed Martin Aeronautics, Marietta, GA 30063While many opportunities exist to leverage additive manufacturing technology for design improvement, technologies must demonstrate reliability and integrity sufficient to satisfy certification authorities in order to open the door for use on flight vehicles. This certification challenge is formidable given the large number of process parameters, the magnitude of process variability, and the sensitivity of mechanical properties to these process variables.

This certification challenge is formidable given the large number of process parameters, the magnitude of process variability, and the sensitivity of mechanical properties to these process variables. Historical precedent in certification of other materials and processes, such as composites, castings, and welded joints that exhibit sensitivity to parameters, geometry, and operator skill provides a guide to certification of additive structures. This paper discusses certification processes for these components and applies lessons and methods from them to develop a strategy to certify an additively manufactured structural fitting component.structural componentfabrication with additiveI. IntroductionAdditive manufacturing (AM) has the potential to revolutionize design of many types of structural components if the challenge of designing and fabricating components with reliability sufficient for certification can be met.

Opportunities to leverage AM processes for structural improvement include low volume production, aging component replacement, piece part reduction, multi-functionality, increased complexity, and performance improvement through physics-based design optimization free from many conventional manufacturing constraints.Despite these opportunities, the certification challenge is formidable and must address reliability throughout the component life cycle through material integrity, manufacturing process integrity, design and analysis method integrity, and robust testing and inspection built on a foundation of statistics. Knowledge, methods, and tools in each of these domains must be correlated and calibrated against robust testing at multiple scales.

Knowledge, methods, and tools in each of these domains must be correlated and calibrated against robust testing at multiple scales. As with other manufacturing processes, a comprehensive building block test program, tailored to AM process specifics, can be effective in building reliability by developing statistical confidence through testing, analysis correlation, and inspection at various levels of structural complexity, starting with small coupon specimens and progressing through structural elements, components and full-scale products. Only through such rigorous, structured investigation of manufacturing capability will reliability of AM structures improve sufficiently to enable widespread application of this promising technology.This need for rigor and structured methodology to ensure reliability must be balanced with the unique demands and opportunities of additive manufacturing.

For example, if additive manufacturing is to be used for low-volume production of a component, execution of a full building block test program with the thousands of coupons typical for composite structures is not economically feasible. Another consideration is the effect of increased geometric complexity on the validity of assumptions used in conventional analysis methods widely used to write margins-of-1 Professor in Practice, University of Texas at Arlington 2 CEO, Titan Industries 3 Principal Engineer, Lockheed Martin Aeronautics© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 1 1985safety for aerospace structures. Still another consideration is the ability of AM materials to equivalently replace wrought materials in a replacement of an aging component that was certified potentially many decades ago.

Still another consideration is the ability of AM materials to equivalently replace wrought materials in a replacement of an aging component that was certified potentially many decades ago. Yet in all of these considerations, the structural reliability requirement stands and must be addressed effectively and economically.Developing methods to ensure reliability in additive structures requires industry and researchers to examine sources of variability and determine methods to demonstrate it is within acceptable limits at all size scales across fabrication processes, material properties, design and analysis methods, and inspection capabilities. Fabrication processes must be demonstrated to be stable and controlled with process specifications that yield consistent component performance and enable supplier qualification. Specifications for monitoring and control of raw materials must ensure material integrity during handling and storage while specifying required environmental control.

Specifications for monitoring and control of raw materials must ensure material integrity during handling and storage while specifying required environmental control. Specifications for monitoring and control of process parameters must call on well-characterized parameter sensitivities and specify variances and control limits to communicate acceptable settings and ranges. With these specifications, processes must be fixed to ensure consistency and repeatability from part-to-part, location-to- location, and supplier-to-supplier.To establish design and manufacturing process reliability, fabricated components must be inspected, tested, and evaluated to verify that as-built properties and part performance matches predictions. Historically, this has been accomplished through a combination of non-destructive inspection (NDI), destructive testing, and part-specific acceptance testing using NDI and traveler or witness material to verify process control.

AM processes challenge these standard methods due to the increased sensitivity to and interaction among process parameters. For example, since AM material properties exhibit sensitivity to geometry (e.g. wall thickness) and interlayer time interval effects due to size effects, witness coupons printed in the same build must be carefully planned to ensure they exhibit similar geometric and build characteristics representative of the fabricated part.Current structural design practice ensures reliability through reliance on well-characterized, statistically-based mechanical properties developed through a robust coupon test program that incorporates manufacturing process capability. These allowable property values ensure designs account for material and process variability within acceptable statistical ranges and confidence levels.

These allowable property values ensure designs account for material and process variability within acceptable statistical ranges and confidence levels. AM challenges current design practice again due to the increased sensitivity to and interaction among process parameters and their interaction with design configuration. Because properties can vary significantly due to size variation, surface roughness, and porosity, all of which exhibit strong sensitivity to process parameters, material allowables must be developed with respect to significant process parameters.Finally, structural reliability requires predictable structural performance determined using accessible, verified, and validated design tools. Many AM material and process characteristics challenge assumptions used in standard analysis methods and tools used to develop margins-of safety for aerospace structural components.

Many AM material and process characteristics challenge assumptions used in standard analysis methods and tools used to develop margins-of safety for aerospace structural components. For example, some components use plastic bending analysis to allow loading to develop stresses in the material plastic range. AM part behavior must be characterized to determine if build orientation effects on anisotropy, surface roughness, and porosity can be accommodated in the underlying assumptions that form the basis of plastic strain redistribution in plastic bending analysis.Of course, industry teams can continue to pursue a part-specific certification process that tests the design, material, and fabrication process against all criteria, loads, and parameters without verified and validated allowables, design methods, and inspection process. This part-specific process is unique to individual parts, expensive, and prohibitive to broader application of AM to realize its potential benefits.

This part-specific process is unique to individual parts, expensive, and prohibitive to broader application of AM to realize its potential benefits. Instead, a lean, tailored building block approach is needed in order to provide a comprehensive, standardized strategy that is accepted by certifying authorities to enable AM as a viable manufacturing option. Indeed, researchers at the United States Air Force Research Lab highlighted the many qualification challenges for additive structures but concluded that despite these challenges, the main barriers to broader application of AM are cost, performance, and schedule.1 A standardized, comprehensive strategy is need to improve the cost, performance, and schedule outlook.This paper discusses certification processes for conventionally manufactured components and applies lessons and methods from them to develop a strategy to certify an additively manufactured structural fitting component.

First, conventional structural certification is discussed as a basis for AM processes against those for wrought metal, composites, casting, and welding. Next, a strategy is proposed and discussed for AM structural certification, including building block for AM, coupon testing and inspection, component testing and inspection, and analysis method development, calibration, and verification. Finally, this AM certification strategy is applied to design of a structural fitting to illustrate implementation and identify current capability gaps.© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 2 1986II. Conventional Structural CertificationA review of certification processes for structures manufactured using conventional (non-AM) manufacturing processes illustrates the various paths that are used to ensure structural reliability in aerospace vehicles.

First, an overview of structural certification processes and requirements is given to establish context. Next, certification process details specific to components fabricated from different conventional processes are given. These processes include wrought metals, composite materials, castings, weldments, and powder metallurgy. In particular, the last four processes highlight how reliability is demonstrated in structures where there is high sensitivity of mechanical properties to process parameters and controls.A. Structural Certification Processes and RequirementsBefore an aircraft model can be sold and delivered for service, it must be certified by government regulators that it meets the government specifications, as applied to the specific aircraft program. Part of the requirements that must be fulfilled for certification is to present proof of compliance to the requirements to the certifying agency.

Part of the requirements that must be fulfilled for certification is to present proof of compliance to the requirements to the certifying agency. This is required for all systems on the aircraft, but this discussion is limited to the aircraft structure. The demonstration artifacts that must be presented for structural integrity to obtain certification include the design drawings, analysis and correlated test results. At the end of this process, if the certifying agency is satisfied, the aircraft is certified safe to fly and can then be sold and flown.Typical certifying agencies are the Federal Aviation Administration for commercial aircraft and the U.S. Air Force, U.S. Navy or U.S. Army, depending on the military aircraft. Each agency has its own regulatory documents governing structural integrity and airworthiness such as 14 CFR Part 25 for Transport Category Aircraft2 for the FAA.

Each agency has its own regulatory documents governing structural integrity and airworthiness such as 14 CFR Part 25 for Transport Category Aircraft2 for the FAA. Additional guidance for FAA-certified aircraft and rotorcraft exists in other parts of the code of federal regulations. The Joint Services Specification Guide (JSSG-2006)3 provides the requirements for the Air Force and Navy. Additional Air Force requirements for the Aircraft Structural Integrity Program (ASIP) are called out in Mil- STD-1530C4 and AWB-10155.Both the airworthiness requirements of 14 CFR Part 25 and the ASIP program require the classification of partsbased on criticality.

Table 1 provides a summary of the parts classification categories.Table 1 Summary of Parts Classification RequirementsMilitary Certification (JSSG-2006 and Mil-STD-1530C) Fracture Critical I or II (FC) – Highly critical, single load path safety of flight parts Durability Critical (DC) or Maintenance Critical (MC) – Structural parts affecting the economic life of the aircraft Normal Controls – All remaining structural partsCivilian Certification (FAR-25, et. al.)Principal structural elements (PSE) and Detail Design Points (DDP) – Structural parts, the failure of which could result in catastrophic failure of the airplane. Also called “Critical Part.”Aircraft typically use strain-based certification.

This is also sometimes referred to as the Test-Model-Analyze-Test Correlation approach and can be described asTest – obtain statistically based basic material properties from a stable, documented process  Model – mathematically model the structure  Analyze – perform analysis on the structure using test-correlated analysis approaches  Test Correlation – test the finished part either separately, as part of a full scale component or as part of a full-scale airframe and correlate the strains to finite element based predictions and standard analysis approaches.This approach is used because of the complexity of the design and the impossibility of testing each part withinthe airframe to failure in a full scale static or fatigue test.If a material and its governing specifications have statistically based material properties published in the Metallic Materials Properties Development and Standardization (MMPDS)6 handbook then all certifying agencies recognize the values and they may be used in analysis without further demonstration of compliance.

Parts made from those materials may still require full-scale testing as part of a larger assembly or if they are of unusual configuration or application. If a material is not published in MMPDS, then the original equipment manufacturer (OEM), material supplier, or both must undertake the development of the statistically based material properties necessary as a part of the certification process.7 If a material is process sensitive, such as an additive manufactured material, then the certifying agencies have indicated there will be additional data requirements to obtain certification. Although the certifying agencies have not yet defined these additional requirements, an approach is provided in this document that© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 3 1987is based on historical approaches used as other new materials and processes have been introduced into aircraft design.B.

Conventional Wrought Metal Structure CertificationFor conventional wrought material product such as sheet and plate materials the normal design cycle would be to design and analyze the part using statistically based material properties obtained from MMPDS6, if available. For a new conventional wrought material, these material properties would need to be developed via a battery of tests as described in Section 9 of MMPDS, which typically involves multiple heats and lots of materials in order to capture the process variabilities which might be present. If there is more than one manufacturer of the material, data is obtained from each manufacturer. The analysis performed on these parts, whether based on finite element analysis, classical theory, or a combination of the two, is well understood.

The analysis performed on these parts, whether based on finite element analysis, classical theory, or a combination of the two, is well understood. What is often forgotten, however, is that in the early days of metal aircraft design, there was a significant body of part and component testing performed by individual aircraft manufacturers and by entities such as the National Advisory Committee for Aeronautics (NACA) and universities to develop the confidence that analysis methods using the fundamental material properties adequately and accurately represented part behavior. It is that body of work that provides the confidence that for most new wrought materials there is a straightforward path to certification.Forgings and extrusions provided a variation to the certification requirements for wrought plate. Although these product forms are also wrought product, they do have some dependencies which are a function of the forging or extrusion dies and final product shape.

Although these product forms are also wrought product, they do have some dependencies which are a function of the forging or extrusion dies and final product shape. Typically for forged or extruded product forms, typical shapes are manufactured, cut up and tested in the same way as plate material. Care must be taken to obtain test specimens from all grain orientations within these sample parts. From these sample part tests, statistically based material properties are developed which are then used in preliminary design. Once the specific parts are designed and manufactured, additional testing is required. The extent of the testing is a function of criticality of the part and how different the part is from the sample test parts. If the designed part is a single load path fracture critical part, then sufficient testing to result in part-specific statistically based material properties may be performed.

If the designed part is a single load path fracture critical part, then sufficient testing to result in part-specific statistically based material properties may be performed. Additionally, excess material will be added to each part manufactured which can be made into prolongation coupons. The prolongation coupons may be tested for every part or for each lot of parts to ensure the integrity of the as-manufactured part. Typically the prolongations are tensile ultimate and yield and possibly fracture toughness specimens. The same testing and development effort that verified the analysis methods used adequately predicted the behavior of wrought plate and sheet also included forgings and extrusions.C. Composite Structure CertificationDesign criteria for composite structure include a wide range of possible failure modes, environments, and economic considerations.

Composite Structure CertificationDesign criteria for composite structure include a wide range of possible failure modes, environments, and economic considerations. Depending on the mission and application, these criteria can include: design loads, fastened connections, bonded connections, static strength, stiffness, durability, damage tolerance, crashworthiness, producibility, and maintainability. Each program delineates specific requirements for composite structural design criteria in accordance with the guidance from the applicable certifying agency as described in Section II.A above. Like AM materials, composite materials exhibit significant process sensitivity because the material only achieves a usable state as part of the component fabrication process.

Like AM materials, composite materials exhibit significant process sensitivity because the material only achieves a usable state as part of the component fabrication process. Consequently, appropriate material system allowables, design and analysis methods, manufacturing processes, and inspection processes must be defined, implemented and controlled in order to ensure the entire process produces robust structures that repeatedly satisfy the requirements. As AM is experiencing now in early applications, safe application of composites to primary structure initially progressed slowly as experience and methods for testing, inspection, and analysis evolved.8In order to maximize effectiveness and simplicity in achieving required reliability given this process sensitivity, aircraft manufacturers follow a building block approach, which relies on analysis supported by test evidence, as the principal certification methodology for composite structures.

As shown in Figure 1 below, the building block approach9 conducts analysis and associated tests at increasing levels of structural complexity, beginning with small coupons and progressively increasing complexity through elements, sub-components, components, and full scale product. Specimen count is large (1000’s of coupons) for the simple specimens and decreases as complexity increases.10© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 4 1988Figure 1—Building Block ApproachOnly laminate level, non-interactive criteria are used for margin-of-safety calculations of composite structures.11 The nonhomogeneous nature of composite materials complicates prediction of the onset of failure beyond acceptable levels of probability and economy, thus precluding the use of micromechanical approaches based on constituent material and interface behavior.

Likewise, lamina level failure criteria based on experimental data from unidirectional laminae or laminates that combine the effects of several different failure mechanisms into interaction formulas have been unable to provide sufficient generality and robustness to capture complex interactions between laminae and are also not used. Consequently, statistical basis allowables for composite material systems are developed at the laminate level from fiber-dominated laminate families that encompass the entire design space.11In addition to capturing complex interactions, this laminate level approach captures manufacturing and inspection process capabilities.

Test articles are fabricated using production process specifications and coupons are included that demonstrate the effect of acceptable manufacturing defects at a detectable flaw size.Integral to the building block approach is structural testing that supports development, calibration, verification, and validation of the structural analysis methods that will be used to write margins-of-safety at all critical locations in the structure. This robust connection between analysis and test establishes confidence in the accuracy of the analysis predictions.D. Cast Structure CertificationCast metallic materials bring additional challenges to the certification process. Because the final material and the part are formed at the same time, castings do fall into the category of process-sensitive materials and can have a high degree of process variability. For most aluminum castings included in MMPDS6, only specification minimums are published.

For most aluminum castings included in MMPDS6, only specification minimums are published. The result is that aluminum castings are almost never used in normal aircraft structural applications and then only for non-critical structure. More often they are found in fuel or control systems. Hot Isostatic Press (HIP) and annealed titanium castings do find their way into structural applications and there are statistically based material properties published for this material and product form. The HIP process removes some of the variability thus providing a more reliable part.The certifying agencies require additional casting safety factors as a part of the certification process and, if a casting is used in a structural application, there is additional testing required based on the criticality of the part. Table 2 summarizes these factors and additional inspection requirements for critical castings based on the selection of casting safety factor for commercial aircraft.

Additional guidance is available for non-critical castings from Reference 2 and for military hardware from Reference 3.© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 5 1989Table 2 Sample of Casting Safety Factor and Inspection Requirements per CFR 14 Part 25.621TypeCasting Safety FactorDiscussionEach foundry and process variation has a material property coefficient of variation (COV) equivalent to wrought alloy, similar compositionCritical Casting1≥1.00Prolongations and periodic cut-up required  100% Surface inspection (visual + die penetrant); Radiographic inspection of Structurally Significant internal areas and areas of likely defect1 casting – static test (Strength and deformation)  100% Surface inspection (visual + die penetrant); Radiographic inspection of Structurally Significant internal areas and areas of likely defect1 casting – static test (Strength and deformation)  100% Surface inspection (visual + die penetrant); Radiographic inspection of Structurally Significant internal areas and areas of likely defect3 casting – static test (Strength and deformation)  100% Surface inspection (visual + die penetrant); Radiographic inspection of Structurally Significant internal areas and areas of likely defect3 casting – static test (Strength and deformation)  100% Surface inspection (visual + die penetrant); Radiographic inspection of Structurally Significant internal areas and areas of likely defect1 casting – static test (Strength and deformation)1Failure could preclude continued safe flight and landing of the airplane or could result in serious injury tooccupantsThe selection of casting factor requires the typical flaws and defects found in cast parts are well characterized and sufficient inspection studies have been performed to have a high confidence in the probability of detection of these flaws.

This is crucial in setting the certification requirements for castings and will also play an important role in the certification of process sensitive AM materials.Finally, full scale testing of cast parts for correlation to structural analysis indicates that, for the most part, castings behave in a similar fashion to wrought materials and the familiar classical analysis methods do apply except that typically a reduction is required for the low elongation present in castings. Additionally, the use of the plastic bending analysis techniques are not recommended as they can be non-conservative for castings.E. Powder Metal Structure CertificationPowder metal (PM) components present similar challenges and are treated similarly to cast components. The presence of voids and porosity in PM components is a fundamental characteristic and void size and spacing depends on density developed during pressing as well as post-process heat treatment.

The presence of voids and porosity in PM components is a fundamental characteristic and void size and spacing depends on density developed during pressing as well as post-process heat treatment. Additionally, powder material state must be assessed in terms of powder size, shape factors, and agglomerates as these also contribute to porosity levels. With these characteristics, PM components have testing, inspection, and analysis factors applied in a similar manner as cast components, although the magnitude and scope of these additional processes varies depending on the severity of the porosity developed within the PM process parameters.F. Welded Structure CertificationWelding is another process that exhibits sensitivity in quality and achievable mechanical properties to process parameters and operator skill.

Welded Structure CertificationWelding is another process that exhibits sensitivity in quality and achievable mechanical properties to process parameters and operator skill. The American Welding Society has developed a multi-pronged approach to standards to qualify operator skills and specify procedures for specific geometries and parts based on test results. At the top level, fusion welds are qualified based on a Welding Procedure Specification (WPS), as given in AWS B02.112, which specifies part deposition procedures relevant to a specific application. A WPS details the welding variables required for a specific application to ensure repeatability when the weld is performed by a qualified welder. The WPS cites Procedure Qualification Records (PQR) for specific deposit geometries that are tested to ensure good weld performance according to specific industry standards.

The WPS cites Procedure Qualification Records (PQR) for specific deposit geometries that are tested to ensure good weld performance according to specific industry standards. For the aerospace industry, AWS D17.113 contains requirements for fusion welding of aerospace hardware, specifies qualification requirements for welders, welding operators, and welding procedures, and gives requirements for design, fabrication, inspection, and repair of aerospace welded connections. Finally, welders must show knowledge and skill for a specific WPS through individual certification. This is accomplished with a qualification test that is documented in a Welder Qualification Test Record (WQTR) as defined in the industry standard.This multi-pronged approach to welding qualification has relevance to a general approach to AM certification procedures. It highlights the interdependence of geometry-, process-, and operator-specific quality of the operation.

It highlights the interdependence of geometry-, process-, and operator-specific quality of the operation. Korbin, et al,1 from AFRL, advocate development of the PQR/WPS approach to AM qualification procedures by© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 6 1990generating design databases for specific AM deposit features (the PQR) that can be used to substantiate specific part deposition procedures.The small amount of weld allowables in MMPDS6 are given as a reduction factor on the parent material strength. This is acceptable because welds are used for joining and typically make up a small percentage of the overall part. In addition, another 15% joint efficiency factor is added for analysis for some welded materials. These factors are typically conservative. Finally the certifiers require full scale testing, typically both static and fatigue and the use of the welding procedure specification discussed above.III.

Finally the certifiers require full scale testing, typically both static and fatigue and the use of the welding procedure specification discussed above.III. Proposed Additively Manufactured Structural CertificationAny additively manufactured structural certification approach will need to include elements from the various certification methods described in Section II, including some form of a building block approach, testing and inspection at multiple scales, analysis method development and validation, and standards development. Much of this effort is well underway across a wide range of organizations.A. Building Block Approach for AMAM structures must follow a building block approach tailored to AM materials and processes. While evolution of this process is in its early stages, its scope and characteristics can be deduced from the knowledge and lessons embedded in the certification approaches discussed above.

Some pertinent observations follow.The test-model-analyze-test correlation embodied in building block approach testing provides confidence that analysis methods accurately represent material behavior.A qualified manufacturing process is required that produces controlled, consistent materials so that structural design and analysis uses statistical basis material properties and allowables.As composite structures exhibit high variation at the micro scale, inhomogeneity at the meso scale, and interactions among constituent materials that are difficult to predict at these scales, their properties are developed and applied at the macro scale in order to embed interactions and manufacturing effects at a scale where they can be effectively characterized.The process sensitivity that occurs in cast structures is effectively accounted through mitigating post- process heat treatment that reduces variability and use in analysis of casting factors dependent on the nature and scope of additional testing and inspection requirements.

For critical applications, this can include part- specific manufacturing process qualification.The mechanical property degradation (i.e. low elongation, or ductility) seen in castings drives allowable reductions and precludes use of certain analysis methods (i.e. plastic bending analysis).The geometric sensitivity seen in welded structures is effectively accounted through geometry specific process qualification.Porosity levels in powder metal components depends on the powder size, shape factors, and agglomerates of the powder material state.Robust, comprehensive standards are the foundation for repeatable, reliable manufacturing processes, testing, inspection, and analysis.From these observations a framework is proposed for an AM building block approach that accounts for variability at multiple scales and is sensitive to process parameters, material characteristics, part geometry, and interrelationships among these variables.

Development of statistically based allowables is required, but low volume production (a key AM advantage in some applications) challenges the economics of large coupon test programs. These material properties will be associated with specific material and process specifications, and depending on the process variabilities, perhaps to specific manufacturing machines. Additionally key features or design details of some parts may require element level testing for determination of notch sensitivity or detail robustness. Finally full scale component testing will be required to determine if there are any additional effects due to scale up of the process, details or build geometry and limitations.B. Testing and Inspection for AMAs part of the building block approach, testing and inspection occurs at all size scales.10 The nature and scope ofsuch testing in the context of AM is discussed below.1.

Material and Process Specifications and TestingMaterial specifications for the raw material which define constituents, raw material processing, storage, handling and quality control are required to define the material to be used in the manufacture of aircraft parts. The process© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 7 1991specifications should be established once the process is reasonably well understood and required process controls codified in the process specifications. Process specifications must fully define the manufacturing process space.

thermal treatments, surface finishing, etc)Quality Assurance and Acceptance Requirements In-situ inspection requirementso o Visual (part) o NDT (part) e.g., inspect: Porosity, discontinuities, cracks, incomplete fusion and acceptance limits o Mechanical (process control specimens) e.g., test: tensile, fracture toughness, as required, surfaceroughness, as required and acceptance limitso Metallurgical and Composition (process control specimens) – Inspect: microstructure, minimumand maximum elemental limits2. Coupon Testing and Statistically Based Material PropertiesCoupon level testing is needed for material characterization and for basic material properties for use in the design process. It is important that the variability be captured in the generation of the material properties, including machine to machine variability. Table 36,14 provides the static strength subset of a proposed test matrix based on the requirements Chapter 9 of Reference 6.

Table 36,14 provides the static strength subset of a proposed test matrix based on the requirements Chapter 9 of Reference 6. In addition, to the testing for static strength properties, testing for fatigue and fracture properties would also be required. The test coupons would be taken from an engineered test part similar in size and shape to anticipated usage which allows for coupons to be obtained for all material directions. To capture the process variability, multiple heats and lots and multiple machines should be used in the generation of this data. From Reference 14, “a heat is defined as a single uniform composition. For a powder based process a heat may be defined as a heat or lot based on the AMS7001 definition. For a wire based process a heat is based on a single wire lot or lots of the same size produced from the same ingot in consecutive (continuous) runs.” Builds are separate processing of parts where feedstock, process parameters, configuration or machine have changed.

Prior to the start of testing, the material and process specifications must be stable and mature. At the current state of the art, the material properties generated from Table 3 could only be used for preliminary design to initially size the part, followed by significant part-specific material property testing or full scale point design testing. As a result, an argument could be made for a reduced test matrix to obtain only preliminary design properties for use in initial sizing.

However, it is anticipated that as more information becomes available and the processes become more fully characterized that there will be a path for the use of statistically based properties for design and analysis of additively manufactured parts based on Table 3, much in the same way as is the current practice for wrought, cast parts or composite parts.© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 8 1992Table 3 Typical Test Matrix for Statistically Based Static Material Properties14Test TypeAddl InfoX-OrientationY- OrientationZ-OrientationTypeHeatsLotsBuildsStatistical BasisFtu100100100Direct101010A and BFty100100100Direct101010A and BFcy202020Derived31010A and BFsu202020Derived31010A and BBearing Yielde/D=1.520 (X-Y)20 (X-Z)Derived31010A and Be/D=2.020 (X-Y)20 (X-Z)Derived31010A and BBearing Ultimatee/d=1.520 (X-Y)20 (X-Z)Derived31010A and Be/D=2.020 (X-Y)20 (X-Z)Derived31010A and BCoefficient of Thermal Expansion1010TypicalElastic Modulus, Tension9993Multiple3TypicalElastic Modulus, Compression9993Multiple3TypicalStress-Strain Curves to Yield66366TypicalStress-Strain Curves full Range66366TypicalDensity1010Elongation303030333S-BasisHeat/Lots or Builds, as applicableDuplicate Measurements3.Inspection MethodsAs discussed in item 1, inspection requirements and accept/reject limits are a part of the process specification; however, additively manufactured parts provide both some opportunities for in-situ inspection15 that are not available for other material product forms as well as some challenges relative to final part inspection.

These methods do not offer sufficient fidelity, as yet, to replace post-manufacture quality inspections. In-situ monitoring is most useful in ensuring the process stays within specification parameters. Post-manufacture inspection ensures the final, finished part meets all of the requirements and may take the form of non-destructive inspection using radiographic or ultrasonic inspection techniques, part cut-up and metallurgical examination or the testing of traveler or prolongation coupons.C. Analysis Method Development for AMWhile individual components can be certified by test in the absence of material allowables and validated analysis methods, this approach does not scale to a general certification approach for AM due to the large expense of testing every geometry, loading, environment, and failure mode for each individual component.

Nevertheless, this approach must currently be used for AM components given the uncertainties in AM materials and process.A more general and cost- and schedule-effective approach would satisfy structural requirements primarily by analysis supported by testing. For this approach to be a feasible AM structural certification process, validated analysis methods are key. Given the variability in as-built AM materials and the sensitivity to microscale flaws, analysis methods must be closely calibrated and correlated using suitable test data and a controlled manufacturing process. From this, test data must be used to determine appropriate safety factors, based on probabilities of failure for measured process variability levels. Furthermore, appropriate modeling idealizations at appropriate size scales must be correlated to demonstrate analysis methods suitably predict mechanical behavior and critical failure modes and loads.

Aircraft development programs effectively use analysis methods based on classical theories, finite element analysis, and empirical curve fitting or calibration16 to ensure close correlation and failure prediction. Building block tests at coupon, element, component and full-scale aircraft all serve to calibrate and correlate analysis methods.D. Standards for AMThe development of industry accepted standards is a critical step in transitioning additive manufacturing into practice. The creation of standards accelerates the adoption of new technology by allowing for the qualification of new material and processes, specification of requirements, and definition of test methods and protocols. Practitioners in industry, academia, and the regulatory agencies develop standards to communicate guidance, share best practices, and document technical data.

Practitioners in industry, academia, and the regulatory agencies develop standards to communicate guidance, share best practices, and document technical data. Standards are necessary for the additive manufacturing industry in areas such as the following:  Powder/Feedstock  Machines & Operators  RepairRecycling Materials  In-process Sensing  SafetyTechnical Data Packages  Post Processing  Modeling & Simulation© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 9 1993Many organizations are currently involved in generating a set of standards for additive manufacturing. Due to variation between additive processes, standards are process and material dependent. Several challenges are created with different organizations undertaking standards development.

Due to variation between additive processes, standards are process and material dependent. Several challenges are created with different organizations undertaking standards development. These challenges include a high risk of duplication of efforts and the creation of overlapping content, the potential for inconsistencies or even contradictions, and the creation of ambiguity and confusion due to conflicting standards. History has provided many examples of the difficulty and time consuming effort to harmonize standards. The additive manufacturing industry has an opportunity to avoid this challenge. In Table 4 there are multiple collaborative efforts listed.

Draft in work: “Specification for Fabrication of Metal Components using Additive Manufacturing”ASTM International (in cooperation with ISO Technical Committee 261)ASTM F42 Committee and Subcommittees: F42.01 Test Methods F42.04 Design F42.05 Materials and Processes F42.06 Environment, Health, & Safety F42.90 Executive F42.91 Terminology F42.94 Strategic Planning F42.95 US TAG to ISO TC 261Fast track processing of existing ASTM standards Adoption of ISO standards by ASTM Co-development of joint ISO/ASTM standardsASTM InternationalInternational Standards Organization (ISO)American Society of Mechanical Engineers (ASME)ASTM E07.10 (Non Destructive Test Methods)Technical Committee 261Y14.46 SubcommitteeCollaboration with ASTM F42 Standard Guide for NDT of AM Metal Parts Used in Aerospace Applications—leverages NIST Measurement Science Roadmap for Metal-Based AM Air Force Research Labs AFRL-RX-WP-TR-2014- 0162 NASA Engineering and Quality Standard for AM Spaceflight Hardware, EM20 NASA Nondestructive Evaluation of AM, TM-2014- 218560 AM standardization: processes, terms and definitions, process chains (Hard-and Software), test procedures, quality parameters, supply agreements, and fundamentals Product definition practices for AM Implementation of GD&T© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 10 1994In addition to the organizations listed in Table 4, other agencies are also actively involved in the development of standards, either directly or indirectly, including the National Institute of Standards and Technology (NIST), NADCAP, MMPDS and NASA.

NIST is primarily focused on four research thrusts, characterization of AM materials, real-time control of AM processes, qualification of am material, processes, and parts, and AM systems integration. The MMPDS Emerging Technology Working Group for new metallic materials is heavily involved with SAE AMS-AM committee to generate material property data. NASA has developed the draft standard EM20, Engineering and Qualification Standard for Additively Manufactured Parts for Spaceflight Hardware.The regulators are reaching out to subject matter experts to guide their certification criteria. Regulators are looking for standardization across producers through industry specifications.

Regulators are looking for standardization across producers through industry specifications. The FAA has identified the following top five challenges for additive manufacturing: limited understanding of acceptable ranges of variation for key manufacturing parameters, limited understanding of key failure mechanisms and material anomalies, lack of industry databases/allowables, development of capable NDI methods, lack of industry specifications and standardization. The FAA is also concerned with the lack of a robust powder supply base, the chance of an OEM- proprietary technology path, and the low barrier to entry of new, inexperienced suppliers. It is expected that AM will rapidly expand into aviation with increased levels of AM part criticality.

It is expected that AM will rapidly expand into aviation with increased levels of AM part criticality. This highlights the need to develop procedures (similar approach to CMH-17 for composite materials) which would allow individual manufactures to demonstrate that mechanical properties of parts produced at their facilities achieve ‘equivalent’ properties to published statistically-based values.IV. Electron Beam Melting Process ParametersA repeatable process is critical to certification of parts fabricated by additive manufacturing. In contrast to traditional manufacturing processes, a large number of variables are present during additive manufacturing processes. To certify components fabricated with AM, these variables must be observed and their uncertainty quantified.

To certify components fabricated with AM, these variables must be observed and their uncertainty quantified. These may be measured with proper processes and quality procedures.Due to the wide variety of additive manufacturing processes, and differences in process variables between them, a single process (Electron Beam Melting, Arcam Q20) has been selected for the focus of this paper. Similarities in process variables will exist, but ultimately a thorough understanding of each process is required to observe and quantify sources of variation. The Electron Beam Melting (EBM) process is a powder bed fusion method for additive manufacturing that utilizes an electron beam as the thermal source.A brief overview of the Electron Beam Melting process is included below, followed by a discussion of the primary factors that can influence part quality. The factors are grouped in four categories: Powder Quality, Hardware Setup, Thermal Source, and Build Instructions.

The factors are grouped in four categories: Powder Quality, Hardware Setup, Thermal Source, and Build Instructions. Factors to consider in each group are described in Table 5. An understanding of the parameters that can vary during the process will allow for a certification plan that addresses and quantifies each source of variation.A. Electron Beam Melting Additive ManufacturingThe Electron Beam Melting (EBM) process is similar to other powder bed fusion additive manufacturing processes. A thin layer of powdered material in a powder bed is melted by a thermal source. The differences in capabilities and limitations between laser based powder bed fusion processes and electron beam melting (EBM) stem from the choice of thermal source. The EBM system must occur in a vacuum since gaseous atoms will deflect the beam. Also, since electrons are being used, magnetic lensing controls the beam. A conceptual schematic of an EBM machine is shown in Figure 2.

Also, since electrons are being used, magnetic lensing controls the beam. A conceptual schematic of an EBM machine is shown in Figure 2. Similar to that of a scanning electron microscope, electrons are emitted by a crystal cathode that is heated by an applied current. The electron beam is controlled by three magnetic coils, which are housed in the lower column. The first magnetic lens corrects for astigmatism and generates a circular electron beam with a Gaussian energy distribution. The second magnetic lens is used to focus the beam to the desired diameter. The third magnetic lens is used to deflect the focused beam to the desired point on the build platform.Magnetic control of the electron beam allows for beam deflections without the involvement of moving mechanical parts.

This results in high to keep multiple melt pools alive scan speeds andthe abilityFigure 2: EBM Conceptual Schematic© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 11 1995simultaneously, which in turn allows for high build rates. Magnetic control also enables wide beam spots to scan at high speeds to preheat each successive powder layer. This preheating reduces thermal gradients, which minimizes residual stresses as the part is built.Metal powder (45 –106 µm diameter) is fed to the build table from two powder hoppers. A mechanical rake spreads the powder from the hoppers into a thin layer (90 µm) across the build table. Initially, the beam scans at a high speed in multiple passes to preheat powder to a sintered state. After preheating, a slower beam scan is used during the melting cycle. Upon completion of the melting cycle, the build table is lowered by one layer thickness into the build tank and the process is repeated.

Upon completion of the melting cycle, the build table is lowered by one layer thickness into the build tank and the process is repeated. The build platform moves down the z-axis as the build progresses. When all layers have been completed, the built part is allowed to cool inside the chamber, which is then filled up with helium to assist cooling.During the process, the entire EBM build system, build tank and electron beam column is under vacuum. This is required since electrons interact with the gaseous atoms, which may deflect the electrons if present. The use of vacuum is advantageous for the process, since the vacuum will prevent reactions between reactive metals, e.g. titanium, with atmospheric gases, such as oxygen. Additionally, the vacuum also acts an insulator to help keep the process at an elevated temperature.

titanium, with atmospheric gases, such as oxygen. Additionally, the vacuum also acts an insulator to help keep the process at an elevated temperature. The elevated temperature of the build chamber is achievable due to the high power of the electron beam (up to 3 kW) and the ability to rapidly scan the beam in a defocused mode. The elevated build temperature minimizes thermal gradients between the powder bed and part, which helps to prevent residual stress build up in the finished part. The elevated temperature will also lightly sinter the powder particles surrounding the fully melted finished part.B. Factors Affecting Material VariabilityAs discussed in the previous section, the EBM process creates fully dense metal parts through layer by layer melting of metal powders with an electron beam. Material properties of the fabricated parts depend on the thermal history of the build.

Material properties of the fabricated parts depend on the thermal history of the build. The time history of spatially deposited energy affects cooling rates, which in turn affects microstructure. The microstructure of the part, along with any defects created during the process will determine the material properties. Variability in material properties between builds will be created by changes in the thermal history and process defects. Table 5 describes the factors that can create changes in thermal history. These factors are discussed in detail below.Table 5: Factors to Consider with EBMPowder Quality Chemical Composition Physical CharacteristicsHardwareThermal SourceCalibration / Alignment Cathode Age Maintenance Inert Gas FlowCathode CurrentBuild InstructionsControl Software Process Themes Part Position / Packing Density1. Powder QualityDifferences in the feedstock powder used between builds can affect final material properties.

Powder QualityDifferences in the feedstock powder used between builds can affect final material properties. Chemical and physical characteristics of the powder influence thermal history in two ways: Differences in absorption of thermal energy and differences in ability to spread powder.Chemical composition of the powder is especially important when the feedstock material is an alloy. Melting and evaporation temperatures between various alloying elements differ. Thus, for the same application of energy, differences in chemical composition will affect thermal energy absorption, which in turn affects solidification of microstructure, or differences in defects. Chemical composition may influence the rate of defects by variation of elements with potential to evaporate during the build. For example, when using Ti-6Al-4V for feedstock powder, evaporation of Al is likely to occur due to its relatively low vaporization point.

For example, when using Ti-6Al-4V for feedstock powder, evaporation of Al is likely to occur due to its relatively low vaporization point. Evaporation of elements can create porosity in parts, which will affect material properties.The physical composition of feedstock powders affects the ability for the mechanical raking system to grab the proper amount of powder for a layer and its ability to evenly spread powder across the build area. The physical composition includes the size distribution of the powder particles (PSD), as well as the powder shape. Irregular shapes will impact flowability. Flowability of the powder affects the size and shape of the powder heaps beneath the powder hoppers. Variation in these heaps vary the quantity of powder spread across the plate.

Flowability of the powder affects the size and shape of the powder heaps beneath the powder hoppers. Variation in these heaps vary the quantity of powder spread across the plate. Too little powder will create regions of porosity and too much powder may create layers that are too thick to be melted to the layer below, creating delamination.Variation in chemical and physical characteristics occurs between the individual powder blends received from the material suppliers. This variation is inherent to the manufacturing process and unavoidable. Additional changes to powder chemical composition and physical characteristics may occur to the same powder lot over time. Powder© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 12 1996that is recycled from build to build may experience changes in chemical composition due to pick up of oxygen, evaporation of alloying elements, or due to the introduction of contamination.

The physical characteristics of the powder may also change since a majority of the powder in a build is semi-sintered during the process. This sintering will cause shape distortion of powder particles (less spherical) and introduce satellites (smaller particles bonding to larger particles). Both of these phenomenon cause changes in flowability and powder size distribution.2. HardwareThe maintenance and configuration of equipment can cause variability in final parts. Hardware strongly influences the accuracy and precision of powder layers and the time history distribution of energy. Both of these factors create variation in the thermal history of the build. The three major subsystems influencing this most are the rake, table, and electron magnetic lensing. The rake’s main influence over the repeatability of process stems from the systems only goal being to spread a proper layer of powder.

The rake’s main influence over the repeatability of process stems from the systems only goal being to spread a proper layer of powder. In pursuit of this goal the rake does not spread a layer of powder in a consistent amount of time. It varies with powder heap size and powder flowability. This results in a varying cooling times between layers. The table is a mechanical system that moves in the z-axis and, under ideal conditions, is known to vary +/- 5 microns. The electron magnetic lenses are susceptible to outside influences. Moving metal objects, high power eclectic devices cycling nearby, and other magnets can change magnetic fields shapes. These external variables are not measured by the machine.Calibration and alignment of the electron beam are required to apply the desired beam intensity to the desired location. Variation in the beam intensity and location will create changes in the thermal history of the build.

Variation in the beam intensity and location will create changes in the thermal history of the build. Calibration of the magnetic lenses is required to compensate for the local magnetic field. This calibration process is partially manual and as such, can introduce variability. The astigmatism, focus, and deflection lenses are calibrated, which can lead to variation in beam spot size and spatial accuracy. Alignment of the electron beam is performed prior to each build. This is often performed manually and may also introduce variation into beam spot size.As discussed above, low pressure helium gas is introduced during the process. This gas is used to assist in cooling and removal of electric charge. Variation in the flow of this gas could create variation in the process. Too much gas will create interference with the electron beam, effectively reducing the beam energy reaching the powder bed.

Variation in the flow of this gas could create variation in the process. Too much gas will create interference with the electron beam, effectively reducing the beam energy reaching the powder bed. Too little gas will increase the chances of part swelling or smoke events (repulsion of powder during the process, causes porosity and deflection the electron beam by airborne particles).3. Thermal SourceThe electron beam that constitutes the thermal source in this system is generated from a crystal cathode. The electron production has a number of sources of variability. The operator who installs and positions the cathode in the Wehnelt cylinder (also known as the grid cup), positioning of the anode to the cathode, voltage noise present from the power supply, age of the crystal used, and control loop used for regulation of grid voltage. All of these items form variability in the beam current and in turn the beam power applied to the powder.As the cathode ages, the crystal degrades.

All of these items form variability in the beam current and in turn the beam power applied to the powder.As the cathode ages, the crystal degrades. Degradation of the crystal is known to cause an increase in the beam diameter and may influence symmetry. While the magnetic focus attempts to compensate for this degradation, the control is open loop. Thus the accuracy of actual beam spot diameter to commanded diameter (via the focus offset) varies as the cathode ages.4. Build InstructionsArcam’s additive manufacturing equipment contains a multitude of user adjustable parameters. The process themes and control software are the largest factors in build-to-build variability. Process themes and control software define the instructions that tell the electron beam where to be, when to be there, and how much power should be applied. The operator of the machine is in control of all these variables. Scanning algorithms built into the control software may vary from version to version.

The operator of the machine is in control of all these variables. Scanning algorithms built into the control software may vary from version to version. Operating builds with different versions of the control software will also introduce variability. Position, orientation, and packing density of the parts within the build will influence the final fabricated components. Changes to the build packaging will change the thermal history of the build. For repeatable results, the position, orientation, and packaging density must be consistent from build to build. Even though the largest source of variation may be introduce through the build instructions, these factors are the easiest to track and control. A configuration management system that can manage the build files, process parameters, and software versions is necessary for certification.5.

A configuration management system that can manage the build files, process parameters, and software versions is necessary for certification.5. Build DefectsDue to the negative charge of the electrons that bombard the surface, the powder bed will accumulate charge if inadequate conduction exists between the powder particles and the ground. Accumulation of charge increases the risk of electron beam deflection, which will result in a decrease in beam accuracy and spot spreading. Additionally, if the charge in the powder bed builds high enough, the negative repulsing forces between the powder particles will© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 13 1997become higher than the gravitational and friction forces holding them in place. Repulsion of powder during the process will cause porosity and airborne particles may deflect the electron beam.

Repulsion of powder during the process will cause porosity and airborne particles may deflect the electron beam. Accumulation of electrical charge is minimized by semi-sintering the powder bed to increase the conductivity between the individual powder grains, and by additional of low pressure of inert helium gas to the vacuum chamber during the melting process.Variation in location and frequency of build defects will influence the final material properties. Common defects include porosity, delamination, shrinkage, and swelling. Major defects are typically caused by poor part orientation, packaging, or supporting and are addressed with modification of build instructions during development. In a well- developed build major defects will be corrected and the primary concern is variation in porosity. Many factors influence porosity a list of common events follows:Local accumulation of electric charge causes power particles to repel and creates a gap in the powder bed.

Many factors influence porosity a list of common events follows:Local accumulation of electric charge causes power particles to repel and creates a gap in the powder bed.  An arc trip occurs in the cathode, interrupting the electron beam and keeping powder from melting.  Evaporation of alloyed elements, removing material from the powder bed.V. AM Fitting DesignA metallic structural fitting provides a medium to focus discussion of certification issues. In 2013, the GrabCAD General Electric Jet Engine Bracket Design Challenge17, shown in Figure 3, yielded 638 entries ranging from conventional bracket designs to innovative configurations tailored to 3D printing technologies. Some configurations pushed the limits of manufacturability while others would be challenged to meet structural design criteria.

Some configurations pushed the limits of manufacturability while others would be challenged to meet structural design criteria. Nevertheless, design challenges such as this push the engineers to innovate with AM to realize its potentialFigure 3—GrabCAD GE Jet Engine Bracket Challenge16AM technologies might be considered for such a bracket in 3 scenarios, as shown in Figure 4 from left to right,as follows:Conventional design—e.g. as a replacement part on an aging aircraft for cost reduction  Topology optimized solid member design for weight reduction  Lattice structure design for weight reductionCertification considerations for each these configurations are discussed below.a)b)c)Figure 4—3 possible bracket design where additive manufacturing could be consideredWith each of the evolutions of the part the process to gain certification becomes more complex.

The conventional design of Figure 4a made using an additive manufacturing process as a possible low-cost solution to replace damaged or end-of-life structural fittings incurs all of the material and process qualification steps previously discussed and will necessarily require upsizing or the use of a stronger material such as substituting a titanium AM© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 14 1998part for an aluminum wrought part. For this approach stable materials and processes, characterized through coupon testing is necessary. It is also likely that some testing of the full scale part for critical failure modes is required.

For this approach stable materials and processes, characterized through coupon testing is necessary. It is also likely that some testing of the full scale part for critical failure modes is required. Prior testing of components show correlation to the analysis methods is recommended, although it could be accomplished as part of the specific part certification.In recent years structural topology optimization18 has demonstrated its utility in configuring components for improved structural performance.19,20,21 Commercial finite element-based topology optimization software is now being used to configure components to leverage the capabilities22 and work within the constraints23,24 of additive manufacturing. While topology optimization can help determine efficient structural configurations, the design freedom opened by AM often drives these configurations to complex, intricate geometries that challenge some of the assumptions in classical and empirical analysis methods.

Furthermore, any structural criteria not explicitly constrained in the topology optimization is completely ignored and the optimizer will leverage the omission as it converges on a configuration. A topology-optimized solution, such as that shown in Figure 4b, will require all of the testing of the previous example. In addition, it will likely require additional full scale testing for both static and fatigue to ensure that the strength, life and stiffness of the optimized part is comparable to the original fitting.A lattice-optimized part, shown in Figure 4c, would result in the most complex certification process because the lattice structure is an unknown in terms of effect on part life.

The extremely large number of stress concentrations, unknown material characteristics (particularly at lattice intersection nodes), and unsuitability of most classical analysis methods drive dependence on complex finite element modeling for load and stress concentration characterization makes certification problematic. For the foreseeable future, a lattice design such as is shown in Figure 6 likely requires a certification by test approach, thus increasing certification cost and creating part-specific uncertainty in the certification path.VI. ConclusionThis work has outlined a certification strategy for additively manufactured structural components based on the lessons and knowledge developed over many years of certifying structural components from a range of manufacturing processes.

In particular, components fabricated using processes that exhibit strong sensitivity of mechanical properties to manufacturing process parameters provide guidance in how to approach certification of AM structures. Of particular importance in this certification strategy is an integrated approach—an AM- tailored building block approach—that tightly connects testing, design, analysis, and inspection at all size scales. This integrated approach relies on characterization of variability in as-built materials to develop appropriate safety factors, application of robust standards to ensure repeatability, test-correlated and -calibrated analysis methods to ensure physical relevance, and attention to geometry specific-variability.

24 Zhou, M., Lazarov, B. S., Wang, F., & Sigmund, O., “Minimum Length Scale in Topology Optimization by Geometric Constraints,” Computer Methods in Applied Mechanics and Engineering, 293, 2015:266-282.© 2016 by Robert Taylor, Titan Industries, and Lockheed Martin Corporation (PIRA #AER201607010) 16 2000.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperChallenges during laser powder bed fusion of a near-alphatitanium alloy - Ti-6242Si Sagar Patel1, Mohsen Keshavarz1, Mihaela Vlasea1*1Department of Mechanical and Mechatronics Engineering, University of Waterloo, Waterloo, ON N2L 3G1, CanadaCorresponding author: mihaela.vlasea@uwaterloo.caAbstractTi-6Al-2Sn-4Zr-2Mo-Si (Ti-6242Si) is a near-α phase titanium alloy that has a greater strength up to 565 °C compared to the workhorse Ti-6Al-4V alloy with a typical service temperature of up to 400 °C. While there is a wealth of literature to help understand the laser powder bed fusion (LPBF) of Ti-6Al-4V, only a few research articles about LPBF of Ti-6242Si are available in the open literature.

While there is a wealth of literature to help understand the laser powder bed fusion (LPBF) of Ti-6Al-4V, only a few research articles about LPBF of Ti-6242Si are available in the open literature. In this work, LPBF processing diagrams and temperature prediction models were used to investigate the impact of process parameters such as laser power, scan speed, and beam spot radius on macroscale characteristics of the builds such as density and surface roughness. The use of processing diagrams allowed for exploration of density ranges between 99.55-99.98 %, and surface roughness, Sa, ranges between 8-16 μm in Ti-6242Si processed by LPBF. Cracking in Ti-6242Si manufactured by LPBF is reported for the first time.

Cracking in Ti-6242Si manufactured by LPBF is reported for the first time. Cracking during LPBF of Ti-6242Si was observed to strongly depend upon the predicted melting mode (conduction, transition, and keyhole) for a given set of LPBF process parameters.KEYWORDS: Laser powder bed fusion, selective laser melting, titanium, Ti6242, density, roughness, crackingIntroductionLaser powder bed fusion (LPBF) is a metal additive manufacturing (AM) technology that alongside electron beam melting (EBM) has the highest industrial uptake at the moment in the aerospace, automotive, and biomedical sectors [1]. LPBF enables the manufacturing of near net-shape geometrically complex parts. LPBF allows for optimized designs to be explored for manufacturing, such as topology optimized or loading field driven designs for product lightweighting and customization [2], while also reducing environmental impact through input energy reduction and carbon dioxide emissions [3].

The potential of LPBF to build lightweight components has been shown to improve fuel consumption for aerospace and automotive applications, wherein titanium alloys are commonly used [4,5].While additively manufactured titanium alloys are already used for industrial applications in theaerospace [6–8] and biomedical sectors [9–12], the low number of additive manufactured titanium alloys available reduce their potential applications [5]. From the available titanium alloys, the (Ti64) is known as the “workhorse” titanium alloy represents almost half of the total usage of titanium alloys for aerospace applications [6]. When compared to Ti64, the near- known to have improved mechanical properties, particularly creep behaviour [5], which is useful for applications such as high-pressure compressor parts in aviation engines [13].

Alpha ( commonly used in applications where higher strength and toughness, excellent creep resistance, and stress stability is desired at higher working temperatures up to 600°C when compared to Ti64 that cannot be used beyond a service temperate of 400°C [14]. Ti-6242Si, in particular, has been recommended for high temperature applications of up to 565°C [15]. The nominal chemical composition of Ti-6242Si by weight percent (wt%) is: 6 Al, 2 Sn, 4 Zr, 2 Mo, 0.08 Si, 0.05 N, 0.05 C, 0.0125 H, 0.25 Fe, 0.15 O, and balance Ti [16]. Ti-6242Si has also been proposed to be better suited for certain superconductivity applications when compared to Ti64 [17].alloy Ti-6Al-4V𝛼𝛼 + 𝛽𝛽Ti-6Al-2Sn-4Zr-2Mo-Si (Ti-6242Si) isα) and near-Ti alloys areαα1104A survey of open literature shows that LPBF of Ti-6242Si has been investigated by two groups of researchers till date.

Fan and Yang [18] reported a successful use of LPBF to manufacture Ti-6242Si cubic and tensile samples for the first time. Density values of 95.7%-99.5%, hardness of 410 HV, ultimate tensile strength (UTS) of 1437 MPa, and ductility of 5% were reported for the as-built Ti-6242Si samples. A hardness of 450 HV, UTS of 1510 MPa, and ductility of 1.4% were reported for the Ti-6242Si samples aged at 595 °C for 8 hours. Rieger et al. [5] reported improved mechanical properties of LPBF Ti-6242Si at both room and elevated temperatures (500 °C) when compared to LPBF Ti64 and conventionally cast Ti-6242Si. A non-equilibrium hexagonal martensitic microstructure was reported for both Ti-6242Si and Ti64 in as-built condition as is generally observed due to the high cooling rates (1–40 K/μs) associated with LPBF [19]. While using the same LPBF process parameters, Rieger et al. [5] reported smaller acicular grains in Ti-6242Si compared to Ti64, grain structure in Ti-6242Si.

While using the same LPBF process parameters, Rieger et al. [5] reported smaller acicular grains in Ti-6242Si compared to Ti64, grain structure in Ti-6242Si. Nonetheless, cracking during LPBF which were attributed to the smaller prior of Ti-6242Si was not reported in either of these articles.𝛼𝛼 ’𝛼𝛼 ’𝛽𝛽This work is the first attempt at spanning all three melting modes during LPBF of Ti-6242Si, thereby alloy by LPBF. This work focuses on using normalized processingtesting the limits of processing this near- diagrams developed in prior work [20] to investigate the influence of LPBF process parameters on density, side- skin surface roughness, and cracking during LPBF of Ti-6242Si. Laser power, scan speed, and beam spot radius are the principal LPBF process parameters considered for this work.αMaterials and MethodsTwo sets of samples were manufactured in the same build: cylindrical and cuboid samples.

12 cylindrical Ti6242 coupons with a diameter of 10 mm and a height of 30 mm, and 9 cuboids of 5x5x10 mm3 were manufactured on the reduced build volume (RBV) of a modulated Renishaw AM 400 LPBF system. The cylinders were used to optimize the process parameters of the core of the samples targeting the density, and cuboids were manufactured with the purposes of optimizing process parameters to obtain low side-skin surface roughness. The Ti6242 powders obtained from AP&C were plasma atomized, with a particle size range of 15– 45 μm (D10 of 20 μm, D50 of 34 μm, and D90 of 44 μm). A constant powder layer thickness of 30 μm was used for all samples. The build plate layout of the parts with respect to the gas flow and powder flow is shown in Figure 1.Figure 1.

A constant powder layer thickness of 30 μm was used for all samples. The build plate layout of the parts with respect to the gas flow and powder flow is shown in Figure 1.Figure 1. Illustration of the build plate layout on the reduced build volume of the Renishaw AM 400 for the manufactured Ti-6242Si cylinders and cubic samples showing the powder flow direction and argon gas flow.LPBF process parameters such as laser power, scan speed (by means of exposure time), and beam spotradius are known to be the variables that have the highest impact on driving the presence of conduction, transition, or keyhole melting modes [20]. This motivated the LPBF process parameter selection for the core of1105the 12 cylinders shown in Table 1. A hatch distance equal to the beam diameter was chosen for a given cylinder to minimize the potential for lack of fusion defects.The LPBF process parameters that were varied for the cylinders (for density evaluation) are shown in Table 1.

The point distance was kept constant at 55 μm for all cylinders. The beam spot radius variations were obtained by defocusing the beam to positions above the build as illustrated in Fig. 2 of Ref. [21]. Renishaw AM 400 provides a laser beam wavelength of λ = 1070 nm, and the beam spot radius at the focal point is r0 = 35 μm. The laser beam was defocused to two distances (z) of 3.67 and 5.63 mm above the build plate, to obtain divergent beams of radiuses 50 and 65 μm, respectively, as shown in Table 1. These values were calculated for a Gaussian distribution by equation (1). For both cylinders and cuboids, the scanning strategy included scanning of the hatch volume (core) using the meander scanning strategy, followed by a border scan that involves melting of the edge of each layer of a coupon. The direction of core meander scanning was rotated by 67° between successive layers.

The direction of core meander scanning was rotated by 67° between successive layers. For the cylinders, the border parameters were kept constant with power – 100 W, point distance – 45 μm, and exposure time – 40 μs.r=r 0  1   +  λ z 2 π r o  21/2    Table 1. LPBF process parameters used in the core of the Ti-6242Si cylinders – density evaluation.

LPBF process parameters used in the core of the Ti-6242Si cylinders – density evaluation. Border parameters were kept constant for all cylinders.Sample codePower (W)Exposure time (µs)Beam spot radius (µm)Hatch distance (µm)12253050100215030357033003065130437530651305150503570630050501007375506513082255035709300506513010225505010011300305010012225303570In the cuboids, the core was kept constant at the process parameters used for cylindrical sample 10 fromTable 1 but the parameters for the border scan were varied as shown in Table 2 with a laser power 225 W, point distance 55 μm, exposure time 50 μs, beam radius 50 μm, and hatch distance 100 μm. In these samples, the point distance was kept constant at 55 μm, and the beam spot radius was 35 μm (focused beam) for the borders.Table 2.

In these samples, the point distance was kept constant at 55 μm, and the beam spot radius was 35 μm (focused beam) for the borders.Table 2. LPBF process parameters used in the borders of the Ti-6242Si cuboids – side-skin surface roughness evaluation. The point distance was kept constant at 55 μm, and the beam spot radius was 35 μm (focused beam). Core parameters were kept constant for all cuboids.Sample codePower (W)Exposure time (µs)A100251106(1)B25045C10045D10065E17545F17525G25025H25045I25065X-ray co mputed to mography (XCT) methodsThe manufactured cylinders were removed off the build plate by electrical discharge machining (EDM) and analyzed for porosity characteristics using a 3D X-ray computed tomography (CT) scanner (ZEISS Xradia 520 Versa) using a 11 μm voxel size. The X-ray CT scanning parameters used are shown in Table 3.

The X-ray CT scanning parameters used are shown in Table 3. To visualize the pore distribution within each sample, the CT scans were analyzed using an image processing software (Dragonfly 3.0, Object Research Systems Inc.). The brightness, contrast, and opacity of the data volume were adjusted to highlight the features of interest within the CT scans. For each cylinder, the entire diameter of 10 mm and approximately 8.5 mm of height along the build (Z) direction, along the central region of the cylinders was analyzed. Cylinders 6 and 8 from Table 1 and Figure 2 cracked on the build plate before the EDM cut. Due to the multiple cracks on the samples that were perpendicular to the build direction, it was not possible to evaluate 8.5 mm height for density. A diameter of 10 mm and approximately 7.2 mm of height for cylinders 6 and 8 was hence evaluated.Table 3. X-ray computed tomography parameters used for scanning the Ti-6242Si cylinders.

A diameter of 10 mm and approximately 7.2 mm of height for cylinders 6 and 8 was hence evaluated.Table 3. X-ray computed tomography parameters used for scanning the Ti-6242Si cylinders. Unit [μm] [W] [kV] - - [mm] [mm] [s]Parameter Voxel size Source power X-ray energy Filter X-ray optic Source position Detector position Exposure time Number of projections Binning levelValue 11 10 140 HE2lens 0.4 23.1391 × 120.787 2.010012Microscopy and imaging methodsThe surface topography of the side-skin surfaces of the cuboid samples from Table 2 and Figure 3 was characterized using the Keyence VK-X250, a laser confocal microscope. The image processing and roughness measurements were completed using the microscopes processing software. A scanning area of approximately 700x500 μm2 was scanned on the 4 side-skin surfaces (surface along the build (Z) direction) of each cuboid, using a vertical z-axis resolution of 0.5 μm.

A scanning area of approximately 700x500 μm2 was scanned on the 4 side-skin surfaces (surface along the build (Z) direction) of each cuboid, using a vertical z-axis resolution of 0.5 μm. Sa and Sz are the surface roughness metrics used in this work. Sa is the arithmetical mean height of a surface area which is defined as the absolute value of the height from the average surface. Sz is the sum of the largest peak height value and the largest pit depth value within a scanned area of a given surface. Surface correction for plane tilt was employed for all scans, to accommodate for the tilt cause by the roughness of the resting surface. The roughness is measured near the middle of the sample height1107i.e., approximately halfway through the build height of the sample for all 4 surfaces of each sample. Measurements of Sa and Sz from the 4 surfaces of every cuboid were used to calculate the average and standard deviation values of Sa and Sz shown in Figure 3.

Measurements of Sa and Sz from the 4 surfaces of every cuboid were used to calculate the average and standard deviation values of Sa and Sz shown in Figure 3. This procedure is similar to the one adopted by Patel et al. [22] and Rogalsky et al. [23].Results and DiscussionProcessing diagrams to identify melting modesThe normalized processing diagrams for the Ti-6242Si cylinders shown in Figure 2 and Figure 3 are obtained through the methodology described in previous work [20]. The terms equation (2) and (3), respectively.andin Figure 2 are given by∗∗𝐸𝐸𝑣𝑣E=q * l=AP eff −λ T l ( 2 m tT 0)v=Bvr αFigure 2. Normalized processing diagrams for the twelve Ti-6242Si cylinders used for density evaluation. The density values obtained from XCT data in percentage (%) are provided next to the points on the diagram.

The density values obtained from XCT data in percentage (%) are provided next to the points on the diagram. The cracked samples are shown by an asterisk sign (*) next to their labels and density values.In equation (2),is dimensionless heat input,is the effective laser power [W], A is laser absorptivity,is powder layer thickness [m], ∗is thermal conductivity [W/(m.K)],is melting temperature [K], and is the laser is thermal diffusivity [m2/s]. Values for the Ti6242𝑃𝑃𝑒𝑒𝑒𝑒𝑒𝑒 is beam spot radius [m], and𝐸𝐸is initial (or powder bed) temperature [K]. In equation (3), 𝑙𝑙𝑡𝑡 𝑇𝑇0 beam velocity [m/s], material constants are taken at the room temperature from Welsch et al. [15] and are summarized in Table 4.

In equation (3), 𝑙𝑙𝑡𝑡 𝑇𝑇0 beam velocity [m/s], material constants are taken at the room temperature from Welsch et al. [15] and are summarized in Table 4. 𝑟𝑟𝑏𝑏 α The dimensionless beam power term, equations (4) and (5) respectively [24,25].is dimensionless beam velocity,𝑇𝑇𝑚𝑚λ∗𝑣𝑣𝑣𝑣 , are defined in, and the dimensionless layer thickness term,∗∗𝑞𝑞𝑙𝑙q=AP − λ r T ( B mT 0)1108(2)(3)(4)l*=⋅ 2 r Bl tThe laser absorptivity, A, values chosen for a given set of process parameters in equation (2) are obtainedfrom equation (6), which is an adaptation of the scaling law for laser absorptivity determined by Ye et al. [26] for low reflectivity Ti, Fe, and Ni alloys using the terms derived in [20].A= − 0.70 1 exp −0.66 mq * vπ  The term,, in equation (6) is obtained by replacing the A term in equation (4) with(the minimum , value for titanium alloys is reported as 0.26 by Ye et al.melt absorptivity).

The minimum melt absorptivity, [26] which is used in this work.∗ 𝑞𝑞𝑚𝑚𝐴𝐴𝑚𝑚𝐴𝐴𝑚𝑚Table 4. Thermophysical properties of Ti-6242Si used in this work [15].PropertiesMaterial (Ti-6242Si)Density [kg/m3]4540Thermal conductivity [W/m.K]7Specific heat capacity [J/kg.K]460Melting temperature [K]1978Vaporization temperature [K]3315Figure 3. Normalized processing diagrams for the nine Ti-6242Si cubic samples used for side-skin surface roughness evaluation with the Sa values in µm (left) and Sz values in µm (right).The effective laser power,, for modulated LPBF systems such as the Renishaw AM 400 is given byequation (7).

Additionally, equation (8) is used to obtain the values for laser beam velocity,𝑃𝑃𝑒𝑒𝑒𝑒𝑒𝑒, in this work.𝑣𝑣1109(5)(6)P eff=t⋅ P t e + ted(7)v=tep d +td(8)is the time when the laser is acting on the material (exposure time) [s], and is the time when the laser is off and repositioning to the next exposure point (drill delay time) - assumed toIn equations (1.3) and (8),𝑡𝑡𝑒𝑒be 10 µs for the AM 400 system. In equation (8), 𝑡𝑡𝑑𝑑 distance) [μm]. The dimensionless peak temperature, 𝑝𝑝𝑑𝑑 thresholds is derived in previous work [20]. For low reflectivity materials such as titanium, the threshold ∗ 𝑇𝑇𝑝𝑝 between conduction mode and transition mode is predicted to be when the vaporization depth exceeds 0.5 times beam spot radius used, i.e., temperature) of Ti-6242Si and its melting point.

The term ∗ 𝑇𝑇𝑝𝑝 𝑧𝑧 dimension along the depth of a melt pool, vaporization depth, threshold between transition mode and keyhole mode is predicted to be when the vaporization depth exceeds 0.8 times the beam spot radius used, i.e.,is the distance between adjacent exposure points (point, in the melt pool used for predicting the melting mode, where 1.68 is the ratio of the boiling point (vaporizationatis dimensionless depth given by the ratio of the corresponds to a∗= 0.5 and the beam spot radius,= 1.68 , i.e., a vaporization depth equal to half the beam radius used. Similarly, the.

Similarly, the. The term∗z∗𝑟𝑟𝑏𝑏𝑧𝑧𝑧𝑧 , as shown in Figure 2 and Figure 3 [20].= 0.5dv = 0.5rbat∗ 𝑇𝑇𝑝𝑝Pore space and cracking characteristics= 1.68𝑧𝑧∗= 0.8The normalized processing diagrams for the density study are shown in Figure 2 which help to predict the melting mode for a given set of LPBF process parameters, thereby enabling an understanding of the physical origin of porous defects in the three-dimensional cylinders. Density values for each cylinder were obtained from the XCT data which are shown next to each cylinder’s marker in Figure 2. It is important to note that the density values obtained from XCT are approximations of the true density and a relative assessment of part quality due to the voxel size (11 μm) detection limit. A lower voxel size would help in detecting finer defects and could thereby lead to lower density values for the samples.

A lower voxel size would help in detecting finer defects and could thereby lead to lower density values for the samples. Nonetheless, Ti-6242Si seems to have a wide-processing window for near fully dense parts as cylinders with over 99.90% density are observed in the conduction mode (cylinder 2), transition mode (cylinders 5, 7, 10, 11, and 12), and keyhole mode (cylinder 8). In Figure 4 and Figure 5, defects with aspect ratios above 0.7 were considered as rounded defects and defects below this aspect ratio were considered as irregularly shaped defects. More description of the porous defect visualization can be found in our prior work [21].An orthographic projection of the porous defect space for 6 representative cylinders along with build plate(XY) plane is shown in Figure 4. In the keyhole mode cylinders (#6 and 8), a length of approximately 7.2 mm in the region between the 2 cracks as shown in Figure 6 were used for the porous defect visualization in Figure 4.

In the keyhole mode cylinders (#6 and 8), a length of approximately 7.2 mm in the region between the 2 cracks as shown in Figure 6 were used for the porous defect visualization in Figure 4. All 6 samples in Figure 4 show a higher concentration of porous defects near the sidewalls of the cylinders. These porous defects are caused by the rapid formation and collapse of deep vaporized regions due to the lower laser beam velocity at the turn points of a given layer which is common across most LPBF systems [27]. The collapse of the vaporized regions of the melt pools leads to trapping of atmospheric gases in the solidified parts, which was argon in this work. The processing diagram for the cylinders in Figure 2 shows that all the samples (including the conduction mode samples) lie above the predicted surface vaporization threshold which is given for Ti-6242Si.

Hence the presence of vaporization related defects near the sidewalls of by all cylinders matches the predictions from the processing diagrams. Additionally, most of the defects near the sidewalls of the cylinders seem to have an irregular morphology as shown in Figure 4. It is known that defects which form due to material vaporization instabilities may have both rounded and irregularly-shaped morphologies; and defects due to collapse of the vaporized region in particular are known to have irregularly- shaped morphologies [20,28].at∗ 𝑇𝑇𝑝𝑝∗= 1.68𝑧𝑧= 01110Figure 4. An orthographic projection of the porous defect space (above 4 voxels) in the build plate (XY) plane from the XCT data of conduction mode cylinders (1, 2), transition mode cylinders (5, 10), and keyhole mode cylinders (6, 8).

The percentage values provided next to the cylinder labels are the density values also represented in Figure 2.The absence of large irregular lack of fusion defects in Figure 4 and Figure 5 indicates that the melt poolstitching caused by the selection of hatch distance and powder layer thickness was appropriate for the given sets of process parameters. The processing window for near-fully dense Ti-6242Si parts by LPBF on the Renishaw AM 400 seems to be particularly focused in the transition melting mode as observed by cylinders 5, 7, 10, 11, and 12 with densities over 99.90% in Figure 2. Amongst the 6 cylinders visualized in Figure 4, cylinder 5 seems to have the highest density of 99.97%. A three-dimensional visualization of the porous defects space distribution in cylinder 5 along the build (Z) direction is shown in Figure 5 which confirms the presence of very few defects along only the sidewalls of the cylinder.

The absence of rounded vaporization instability related defects in the two keyhole mode cylinders 6 and 8 in Figure 4 means that the LPBF process parameters selected for the samples were likely not aggressive enough to cause such defects. Similar to Ti64 (Ti64), Ti-6242Si is expected to form a dominant martensitic microstructure in LPBF [5,18]. Literature suggests a broad processing window across the three melting modes for such materials [29–33], which is the outcome that is observed through this evaluation of melting modes on density of LPBF Ti-6242Si. Additionally, since it is not possible to observe melt pools in three-dimensional components made out of martensitic titanium alloys such as Ti64 [34], the use of normalized processing diagrams alongside non-destructive characterization equipment such as XCT is an efficient method to rapidly develop process parameters for similarly martensitic materials such as Ti- 6242Si.1111Figure 5.

A three-dimensional visualization of the porous defect space (above 4 voxels) along the build direction (Z) from the XCT data of cylinder 5 (highest density), along with the density value obtained from the XCT data.While cylinders 6 and 8 had a high density in the XCT results shown in Figure 4, macroscopic cracks were observed in both samples at 2 separate locations perpendicular to the build direction as shown in Figure 6. An optical micrograph of cylinder 8 along the build (Z) direction in Figure 7 confirms the presence of cracking perpendicular to the build direction alongside a near fully dense cylinder otherwise as observed in XCT results of Figure 4. Cylinder 5 (transition mode) which had a density of 99.97% is also visualized through optical micrographs along the XY plane and the Z direction in Figure 7, thereby confirming the presence of close to no porous defects in the core and subsurface regions as shown by XCT results in Figure 4 and Figure 5.

Cracking during LPBF of Ti-6242Si has not been reported in the previous two studies targeting LPBF of Ti-6242Si [5,18]. Observed cracking during LPBF of Ti-6242Si could be potentially attributed to excessive shrinkage and higher residual stress in keyhole mode as reported by Song et al. for LPBF of Ti64 [35]. There is also a possibility of hydrogen induced embrittlement causing cracks in the martensitic microstructure as observed by Silverstein and Eliezer during LPBF of Ti64 [36]. Additionally, subsolidus cracking during the arc welding of Ti-6211, an alpha-beta Ti alloy, was observed to be caused by poor hot ductility of the martensitic alpha-prime (α′) microstructure by Damkroger et al. [37]. Bowden and Starke [38] had previously found that Ti-6211 structures with a larger prior beta grain size were more susceptible to a high-temperature ductility loss than were samples with a smaller prior beta grain size, but Damkroger et al. [37] found the opposite to be true.

[37] found the opposite to be true. These hypothesis for cracking origins during LPBF of Ti-6242Si will be evaluated as part of future work. Additionally, future work would also involve a reprinting of conduction and transition mode coupons to check the repeatability of the successful process parameters from this work in obtaining near fully dense crack free Ti- 6242Si samples.1112Figure 6. Macroscopic cracking (white arrows) observed in cylinders 6 and 8 from Print 1, which are predicted to have keyhole mode core process parameters as per Figure 2. Build (Z) direction is given by the black arrow.Figure 7. Optical micrographs of cylinder 5 (transition mode) and cylinder 8 (keyhole mode) along the XY plane and the build (Z) direction.

Build (Z) direction is given by the black arrow.Figure 7. Optical micrographs of cylinder 5 (transition mode) and cylinder 8 (keyhole mode) along the XY plane and the build (Z) direction. 1113Side-skin surface characteristicsAn evaluation of the effects of LPBF process parameters for the border scan such as laser power and scan speed (by virtue of exposure time) showcases a trend towards reduced surface roughness (Sa and Sz) in the keyhole melting mode in Figure 3. These findings are similar to results by Patel et al. [22] during LPBF of Ti64, where a preliminary investigation into the effects of border LPBF process parameters lead to Sa values of 9.05 ± 0.37 μm with transition/keyhole mode process parameters compared to 17.05 ± 1.59 μm for conduction mode process parameters.

Figure 8 showcases a significantly reduced impact of balling and adhered partially fused powder particles on the side-skin surface of Ti-6242Si cuboids in the transition and keyhole mode LPBF process parameters, when compared to the conduction mode process parameters. The lowest surface roughness (Sa and Sz) values in Figure 3 were observed in the keyhole mode samples which could be associated to the effects of powder denudation caused by the increasing extent of vaporization shown by Matthews et al. [39], leading to lower side-skin surface roughness during LPBF [22], confirmed by Figure 8. It must be noted that studying the effect of core LPBF process parameters or studying the interaction of core and border process parameters on side-skin surface roughness of Ti-6242Si was outside the scope of this work. Care must be taken to understand the extent to which the border parameters in keyhole melting mode may detrimentally contribute to sub-surface porosity and elevated edges.

Care must be taken to understand the extent to which the border parameters in keyhole melting mode may detrimentally contribute to sub-surface porosity and elevated edges. Subsurface porosity will lead to a reduction in the fatigue life of a given part [21], while elevated edges can lead to the powder recoater hitting the edges of a given sample thereby disturbing the flow of powder on the LPBF build plate, and can also lead to recoater blade damage [22].Figure 8. A three-dimensional height map representation of side-skin surface characteristics along with Sa values for sample A (conduction mode), sample G (transition mode), sample H (keyhole mode).ConclusionsThis work involved the use of process diagrams and a temperature prediction for investigating the effect of LPBF process parameters including laser power, scan speed, and beam spot radius on the density and side-skin surface roughness of Ti-6242Si, a near-alpha titanium alloy.

The key findings are summarized below:1.Normalized processing diagrams and X-ray computed tomography show that Ti-6242Si has a broad process window with parts above 99.90% density observed in conduction, transition, and keyhole melting modes of LPBF. The overall density range obtained was 99.55-99.98 %. The highest density parts (up to 99.98 %) are observed in the transition melting mode for Ti-6242Si, with reduced number of defects in both the core and subsurface regions.2.Subsurface porous defects are the main contributor towards porosity across all melting modes for LPBF of Ti-6242Si. These defects are caused by rapid formation and collapse of deep vaporized regions due to the lower laser beam velocity at the turn points of a given layer, which is common across most LPBF1114systems. Simulations in this work predicted vaporization for all of the LPBF process parameters used, leading to the prevalence of subsurface defects in all samples.3.

Simulations in this work predicted vaporization for all of the LPBF process parameters used, leading to the prevalence of subsurface defects in all samples.3. Keyhole mode LPBF of Ti-6242Si leads to macroscopic cracking perpendicular to the build direction, which will be investigated as part of future work.4. Keyhole mode LPBF process parameters for border scans of Ti-6242Si leads to a side-skin surface roughness, Sa, values of less than 10 μm, which is the lowest roughness value when compared to transition and conduction mode LPBF process parameters. This reduction is associated to reduced effects of adhered partially fused powder particles and balling on the side-skin surfaces of Ti-6242Si. AcknowledgmentsThe authors appreciate the funding support received from Federal Economic Development Agency for Southern Ontario (FedDev Ontario, grant # 814654). We acknowledge the generous donation of the powder from AP&C, QC, Canada and for their support in this project.

We acknowledge the generous donation of the powder from AP&C, QC, Canada and for their support in this project. The authors would also like to thank Hamid Azizi and Mehrnaz Salarian from Burloak Technologies for their inputs and help with metallography of the samples. In addition, the authors would like to acknowledge the help of Henry Ma, Jerry Ratthapakdee, James Farnsworth, and Karl Rautenberg in helping with the deployment and characterization of builds at the MSAM Group in the University of Waterloo.References[1] D. Küpper, W. Heising, G. Corman, M. Wolfgang, C. Knizek, V. Lukic, Get ready for industrializedadditive manufacturing, Boston Consulting Group, 2017.[2] A. Salmi, F. Calignano, M. Galati, E. Atzeni, An integrated design methodology for components producedby laser powder bed fusion (L-PBF) process, Virtual and Physical Prototyping. 13 (2018) 191–202.[3] P.C. Priarone, V. Lunetto, E. Atzeni, A.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCHALLENGES IN ASSESSING THE SUSTAINABILITY OF WIRE + ARC ADDITIVEMANUFACTURING FOR LARGE STRUCTURESAnne C.M. Bekkera, Jouke C. Verlindena, Giorgia GalimbertibaFaculty of Industrial Design Engineering, Delft University of Technology, Netherlands bDepartment of Mechanical Engineering, Politecnico di Milano, ItalyAbstractAdditive manufacturing is known as a disruptive technology, in enabling freedom in shape and on-demand production with little human intervention. At present, large-scale digital manufacturing means are being developed, such as Wire & Arc Additive Manufacturing (WAAM). These could be beneficial in aerospace, automotive and construction industries.

These could be beneficial in aerospace, automotive and construction industries. However, while the technology is rapidly developing, little is known on the sustainability aspects. This article explores how such environmental effects could be assessed for a novel technology, and the production of large-scale products by means of a Life Cycle Assessment (LCA). Forerunning results show possible gains in material usage when compared to traditional manufacturing technologies, and in power consumption when compared to different additive manufacturing technologies.

Future research will focus benchmarking WAAM against alternative manufacturing techniques, including green sand casting and CNC milling.IntroductionAdditive manufacturing (AM) or 3D printing is often seen as a disruptive technology (Geraedts, Doubrovski, Verlinden, & Stellingwerff, 2012), which does not only offer production flexibility and customization, but also material and resource efficiency (Huang, Liu, Mokasdar, & Hou, 2013). Larger-scale products (also called mega-scale rapid manufacturing machines) have been considered since the past decade (Buswell, Thorpe, Soar, & Gibb, 2008), in particular for concrete (Lim et al., 2012) and plastics (Love, 2015). Wire and Arc Additive Manufacturing (WAAM) is a technique in which a shape is fabricated by welding layer upon layer with a robotic arm, until a desired three-dimensional shape has been formed (Ding, Pan, Cuiuri, & Li, 2015).

It can span larger areas than other additive manufacturing techniques for metals (Frazier, 2014); while other AM techniques are bound to a predefined bounding box, WAAM can theoretically print objects of any size. While currently in its infancy, a number of academic and commercial institutes are actively working on WAAM applications, e.g. for aerospace, marine, and construction domains (Busachi, Erkoyuncu, Colegrove, Martina, & Ding, 2015). One such example is the attempt to print an eight meters long metal footbridge, depicted in Figure 1 (MX3D, 2015).Due to the early stage of the additive manufacturing, little is known on its sustainability aspects. This article explores how environmental effects could be assessed for the production of large-scale products by means of a Life Cycle Assessment (LCA). It will start by introducing the background of the WAAM process and byFigure 1. Artist impression of 3d printed bridge (source: MX3D).406exposing the process boundaries.

It will start by introducing the background of the WAAM process and byFigure 1. Artist impression of 3d printed bridge (source: MX3D).406exposing the process boundaries. In a subsequent section, the known (life cycle) assessment methodologies are considered. Before conclusions, forerunning results will be reported, in which the object of study is the aforementioned (to be) 3D printed bridge.BackgroundThe WAAM process in this case is based on gas metal arc welding (GMAW), also known as MIG, The WAAM process depicted in Figure 2. This enables metal deposition by creating an arc with DC power, while an inert gas is added to shield oxygen and pollutants from the weld pool. The workflow starts with a CAD model, which is converted to robot paths by a deposition strategy (Busachi et al., 2015). Investigations into mechanical properties show promising results. When porosities are avoided, mechanical properties are good.

Investigations into mechanical properties show promising results. When porosities are avoided, mechanical properties are good. In the case of titanium WAAM, the strength is only approximately 10% less on average than extruded titanium, with a similar ductility (Wang, Williams, Colegrove, & Antonysamy, 2013). Furthermore, fatigue life exceeded extruded titanium in most tested specimen. Measurements of WAAM printed stainless steel are still in development. Current challenges of WAAM involve improving deposition strategies, to prevent peak development at crosslines, and deposition failures (Mehnen, Ding, Lockett, & Kazanas, 2010).Figure 2.

Current challenges of WAAM involve improving deposition strategies, to prevent peak development at crosslines, and deposition failures (Mehnen, Ding, Lockett, & Kazanas, 2010).Figure 2. Schematic of gas metal arc welding.Life Cycle Assessment (LCA) is a method for analyzing the complete life cycle of a product, process or Life Cycle Assessment system, from raw material acquisition to end-of-life treatment, in terms of environmental effects (The International Standards Organisation, 2006).In accordance to ISO standards, there are four phases in an LCA study:a. Goal and scope definition b. c.Inventory analysis (LCI): an overview of the in- and outputs of the system Impact assessment (LCIA): the LCI flows are converted into simpler indicators, to provide more understanding. Steps in this process, depending on the method used, include classification, characterization, normalization and weighting.

Steps in this process, depending on the method used, include classification, characterization, normalization and weighting. Interpretation phase: understanding, discussion and processing of the data collected in the previous steps. This phase is performed in parallel to the rest. Ideally this step includes an uncertainty and sensitivity analysis, to indicate the reliability of the results.d.Goal and scope definitionThe level of detail and the scope of an inventory analysis depend on the goal of the study (The International Goal Standards Organisation, 2006). In the case of MX3D, the main goal is to benchmark WAAM against similar407Figure 3. System Boundary of the WAAM process.techniques (metal additive manufacturing) and/or conventional techniques, and to find out whether it could have environmental benefits in certain areas versus other manufacturing techniques.

Besides the comparison of manufacturing technologies, the goal of the assessment is to compare the environmental impact of the (future) 3D printed bridge to another metal bridge with the same Functional Unit (FU). The FU defines what exactly is being studied and quantifies the function of the system (The International Standards Organisation, 2006). It serves as reference system for inputs and outputs of the system, and enables comparability between alternate systems. The functional unit of the MX3D bridge is defined as: a metal, 3 meters wide bridge spanning 8 meters, with the capacity to support a load of x Newton, following safety requirements, with a lifetime of 20 years. The mentioned numbers in this FU can be subject to change. Of course, this is a limited definition of the intended product, as it also entails aesthetics (Wallsgrove, 1995), safety (Crespo-Minguillón & Casas, 1997), and ergonomics (Nagamachi, 1997).

Of course, this is a limited definition of the intended product, as it also entails aesthetics (Wallsgrove, 1995), safety (Crespo-Minguillón & Casas, 1997), and ergonomics (Nagamachi, 1997). However, in order to make an assessment as objectively as possible, in line with the ISO14040 standard, these aspects will not be taken into account.The system boundaries represent the borders between what is, and what is not taken into account in an System boundaries LCA. Figure 3 displays the system boundaries of the inputs as defined for WAAM in this study. The green blocks represent aspects that can be measured at the WAAM factory. Examples are the power consumption of the exhaust ventilation, and the consumption of welding wire. Other blocks, such as wire drawing, represent data that needs to be obtained by means of databases, scientific papers, and/or suppliers. The system boundary defines the boundary between what will, and will not be taken into account in the assessment.

The system boundary defines the boundary between what will, and will not be taken into account in the assessment. For instance, take the ‘factory’ branch. The power and gas consumption to keep the factory running (for heating and light mainly), is measurable. However, without setting a boundary it will not stop there. The energy, materials, transport and tools necessary for building the factory itself also indirectly contribute to the environmental impact of the factory and therefore to that of WAAM. Consecutively, the tools for creating the materials and the tools themselves also contribute to its impact. As you might start to see, this chain goes on practically infinitely.There are two main reasons for determining the boundary of the system: time (and money) available, and significance (influence on the total impact). The more in detail and depth the assessment is, the more accurate its results will be.

The more in detail and depth the assessment is, the more accurate its results will be. However, every step further down the tree (as in Figure 3) takes an exponentially larger amount408of time. At a certain point, there is not much value in going deeper down the tree for the assessment due to its low significance on the complete impact.For instance, when focusing on the robot welder, its electricity consumption is taken into account. However, the assessment of the manufacturing of the robot itself is disregarded, considering that determining its indicators would take substantially more effort with many unknowns. In addition, we assumed that the production of the robot itself would not have a significant effect (<2%) on the assessment, due to which it can be ignored in the equation (Vogtlander, 2014). This claim is based on the following reasoning: assume the robot is printing stainless steel for 8 hours a day, 260 days a year, for a duration of 5 years.

This claim is based on the following reasoning: assume the robot is printing stainless steel for 8 hours a day, 260 days a year, for a duration of 5 years. Stainless steel is known to have a large eco-impact, while depending on the parameters, a robot prints on average between half a kilo and a kilo per hour. In addition, welding is also known to have quite a high eco-impact. Adding this up over 5 years, it does not seem farfetched to neglect the manufacturing of the robot.Life Cycle InventoryThe basic in- and outputs of the WAAM system are depicted in Figure 4. The quantity of these inputs, particularly that of welding wire, shielding gas and electricity, can be measured easily. The output of welding spatter, or metal waste, is also straightforward. This quantity can be obtained by determining the difference in the weight of the printed part and that of the input material (mass of input welding wire).

This quantity can be obtained by determining the difference in the weight of the printed part and that of the input material (mass of input welding wire). Measuring the output emissions is complex; specialized equipment is necessary for acquiring this data. Without the availability of this equipment, this data can only be approximated by deriving the emissions from existing measurements of welding. Note that this is a basic diagram for the initial assessment of WAAM itself. When assessing products, depending on the goal of a potential comparison, aspects such as finishing, maintenance and recycling should also be taken into account. The result of the Life Cycle Inventory is a list of the material and energy flows in the system, including raw resources and emissions to air. The welding wire for instance is built up from all materials that are inside the alloy, such as silicon and phosphor, plus all components used for and emitted by processing the material into welding wire.

One can imagine that an LCI is usually a long list which provides little understanding. A life cycle impact assessment (LCIA) aids in making sense of all this data.Life Cycle Impact Assessment and its challengesFigure 4. Basic inputs and outputs of the WAAM process.An LCIA converts the input and output flows of a Life Cycle Inventory (LCI) into simpler indicators that are easier to understand. Although there are different methodologies, an LCIA entails of up to four steps as depicted in Figure 5.4091. An assessment starts by classifying all inventory flows into their areas of impact. Commonly used categories are climate change (carbon footprint), natural resource depletion, stratospheric ozone depletion, creation, eutrophication, human toxicity, and aquatic toxicity (Healy, Lorek, & Rodríguez-Labajos, 2013). ozone2. This classification step is followed by characterization: all substances are multiplied by a factor relative to its impact within its category.

ozone2. This classification step is followed by characterization: all substances are multiplied by a factor relative to its impact within its category. The result is a list with impact categories, expressed with a single indicator each. For instance, climate change is expressed in kg CO2 equivalent (kg CO2eq). If CH4 has a characterization factor of 21, the respective impact of 3 kg CH4 would be 63 kg CO2eq.3. Consequently, these results are normalized. This means the characterized data will be compared to (divided by) a reference value. An example of a reference value is the average environmental impact of a European citizen in a year on the respective impact category. Figure 5: Steps of an LCIA.4. The last step is weighting. In weighting, each normalized category indicator is multiplied by a weighing factor, which are determined by value choices, and added to its damage area of effect.

The last step is weighting. In weighting, each normalized category indicator is multiplied by a weighing factor, which are determined by value choices, and added to its damage area of effect. The international standard (The International Standards Organisation, 2006) disproves of this step since no scientific base exists for this weighting, enabling subjectivity and manipulation of results. In case the choice for weighting is made, it is advised to use the default weighting determined by the used methodology to avoid as much subjectivity as possible.Midpoints and endpoints represent impact categories at different stages in the LCIA process. Midpoints and Endpoints Midpoints are problem-oriented. They are the result of the characterization step. Midpoints express the impact onto the different environmental categories, such as acidification and climate change (the problems). True midpoint data is not normalized.

Midpoints express the impact onto the different environmental categories, such as acidification and climate change (the problems). True midpoint data is not normalized. Due to midpoint data having a scientific base, it has a relatively low uncertainty. Endpoints are damage-oriented. They model the environmental impacts in terms of damage to certain areas. For example, in case of ReCiPe1, these categories are human health, ecosystem health, and damage to resources. Where midpoints stop after the characterization phase, endpoints continue with normalization and (non-scientific) weighting. In addition, how exactly midpoints affect the areas of damage is estimated by complex models, containing many assumptions. This causes a relatively high uncertainty for endpoints. A total endpoint score, being the sum of all endpoints, is often used for comparing products.Figure 6 shows the evolution of significant LCIA methodologies through time.

A total endpoint score, being the sum of all endpoints, is often used for comparing products.Figure 6 shows the evolution of significant LCIA methodologies through time. They in particular differ Challenge #1: Diversity in methodologies in their procedural steps, yielding fluctuating LCIA results for equivalent LCI (Pizzol et al., 2011).. Furthermore, classification and characterization is different for each method. While one method might have 18 midpoint categories (ReCiPe) into which LCI data can be classified, another might have 11 (LUCAS) (Budavari et al., 2011). Differences in characterization can occur due to different interpretations of the impact of certain materials.410Figure 6. Development over time of LCA methodologies (Pizzol et al., 2011).In addition, characterization factors of a number of materials are missing in certain methodologies due to lack of knowledge.

Development over time of LCA methodologies (Pizzol et al., 2011).In addition, characterization factors of a number of materials are missing in certain methodologies due to lack of knowledge. Some methodologies have much more substances in their characterization list than others.Variations in normalization occur as well. One methodology might have more outdated reference data than others for instance. Differences in weighting occur mainly due to varying takes on relative importance, and can cause large deviations.Results can be specified in several ways called indicators; varying in depth, communication ease, and Challenge #2: Specifying LCIA results uncertainty. There are multiple categories for expressing the Life Cycle Inventory.

There are multiple categories for expressing the Life Cycle Inventory. They can be prevention-based (eco-costs), problem-oriented (midpoint), or damaged oriented (endpoint), of which the latter two are much more well-known.Prevention-based assessments (Vogtlander, 2014) created a prevention-based impact model, entitled eco-costs. Eco-costs express the environmental burden of a product on basis of the cost of prevention of that burden. It is a relatively new system, and is not widely used at this point in time. This indicator is expressed in euros.Midpoint (problem-oriented) indicators Midpoint indicators communicate the impact on the midpoint categories, e.g. eutrophication, by all substance flows of a system. With every category having a different unit (e.g. kg SO2eq), comparing different categories does not work. This implies difficulties in comparing products using midpoints.

With every category having a different unit (e.g. kg SO2eq), comparing different categories does not work. This implies difficulties in comparing products using midpoints. It is becoming more popular to apply weighting factors to midpoint data in order to gain more understanding of relative contribution. This adds a lot of uncertainty to the data however.Another way midpoints are used is in single issue assessments. A single issue LCIA considers just one midpoint category, such as acidification, and only analyses and/or communicates that aspect. A commonly used issue is climate change (midpoint: carbon footprint). In some methodologies, the three most relevant midpoints are selected. An advantage of this type of indicator is that it is easy to use, and relatively easy to understand and compare. The disadvantage is that it ignores all other pollutants caused by a system.

An advantage of this type of indicator is that it is easy to use, and relatively easy to understand and compare. The disadvantage is that it ignores all other pollutants caused by a system. In addition, issues, or indicators, can be chosen in such a way that a product, which might in fact be more damaging to the environment, appears eco-friendlier.Endpoint(damage-oriented) indicators Endpoint assessments are designed for raising awareness, and can help companies in producing, and potentially people in consuming, less, or in a more environmental friendly way. The disadvantage is that these calculations are often complex, lack transparency, in need of more assumptions compared to midpoints, and contain a higher subjectivity. The reasons endpoints are so often used are ease of communication, decision 411making, and simplicity for the reader.

The reasons endpoints are so often used are ease of communication, decision 411making, and simplicity for the reader. It is much easier to communicate and compare endpoints in order to make a choice in an eco-comparison of two systems; especially when all endpoints are aggregated into a single (total) score.The first main problems with data availability and quality were encountered when setting up the LCIA of Challenge #3: Availability and quality of data WAAM. Information on the eco-impact of manufacturing processes is difficult to find. Some companies make proper LCAs of their products and processes. However, since it is not obligatory to make this data public, it is usually kept confidential. In addition, data on the manufacturing processes within WAAM’s system cannot be found in open databases. Commercial databases and software, such as ecoinvent2 and SimaPro3, are expensive.

In addition, data on the manufacturing processes within WAAM’s system cannot be found in open databases. Commercial databases and software, such as ecoinvent2 and SimaPro3, are expensive. Even there the quality of data cannot be guaranteed.For an initial, draft LCA, data was retrieved from (Vogtländer, 2012). This data is derived from ecoinvent, which is a renowned LCA database. It only provides the final numbers; not what is included and excluded, or basic information such as year of acquirement. With sometimes large unexplained differences between similar techniques, what causes the differences and what values are closest to reality is unknown. Mistakes might also be present in this datasheet. The reason this datasheet was used is that it was the most useful free data available.Recently, access to SimaPro and ecoinvent was attained. SimaPro, with ecoinvent as its main database, offers many possibilities in creating or adjusting processes.

SimaPro, with ecoinvent as its main database, offers many possibilities in creating or adjusting processes. Striking however was that even basic, traditional processes such as green sand casting, were not present in these databases. This means that for benchmarking WAAM, sufficient research needs to be done to, in addition to WAAM, research the comparative techniques.The final entry to this list of challenges is finding or determining comparative structures for benchmarking Challenge #4: LCIA benchmarking a bridge an object. For benchmarking MX3D’s unique bridge, an impact assessment of a comparative bridge is required. For proper benchmarking, this reference bridge should have the same functional unit. It will be difficult to find an assessment of such a bridge. Therefore, the bridge of MX3D will be compared to a simple theoretical construction with the same FU (as described in section ‘Goal and scope definition’).

Therefore, the bridge of MX3D will be compared to a simple theoretical construction with the same FU (as described in section ‘Goal and scope definition’). This will keep subjective elements out of the equation as much as possible. This theoretical construction could for instance be a row of I- bars welded together (with an added handrail if required).Due to the final design of the bridge still being in development combined with the absence of quality data up until recently, the focus of the LCA is currently still on the impact (and price) of printing a kilo of stainless steel with WAAM.Forerunning resultsA start has been made on the impact assessment of WAAM. This section will discuss some forerunning results. Some LCIA data could be retrieved of similar techniques, using the datasheets of (Vogtländer, 2012) described in challenge #3.An initial, draft LCA has been made of printing a kilo of stainless steel with WAAM.

Some impact values of aspects of the process might be quite different in this assessment compared to reality. Especially the impact of the printing itself. Values for standard MIG welding have been used for this purpose. However, the energy consumption of WAAM welding is different compared to average welding. This also has an influence on the welding emissions.Although the available data at the creation of WAAM’s first rough LCA is likely not very accurate, it still provided enough insight to communicate some preliminary results. With relative certainty it can be stated that the main contributor to the environmental impact is stainless steel itself. The second highest impact came from the welding (WAAM). This impact was much higher for instance compared to that of wire drawing: a process for creating the welding wire from billets of stainless steel. Transport had no significant impact.

This impact was much higher for instance compared to that of wire drawing: a process for creating the welding wire from billets of stainless steel. Transport had no significant impact. The raw material of stainless steel itself being the largest contributor to the LCA is a promising result, since WAAM aims for a reduction in material usage.412Measurements of WAAM show an average of approximately 1,84 kW (1,4 for the welder plus 0,44 for Power requirements the robot). This differs significantly from information found on standard Gas Metal Arc Welding (GMAW). A study of (Sproesser, Chang, Pittner, Finkbeiner, & Rethmeier, 2015) compares four different welding techniques by empirical measurements on robot welding apparatus. Energy consumption for gas metal arc welding equals approximately 8 kW.Table 1: Power of WAAM versus GMAW.Manufacturing technique WAAM GMAWPower (kWh)1,84 8This difference might be caused by a combination of factors.

First of all, the material deposition by WAAM might be lower per unit of time compared to standard GMAW. Another contributor could be the difference in welding equipment. Newer welding equipment will be more energy efficient. Different metal/gas combinations can also have an impact on how efficient a weld is made and what energy is required. Finally, welding two cold metals parts together (with perhaps a different welding metal) requires more power compared to welding a layer on top of one still warm layer. This difference in energy consumption should result in a lower impact for WAAM compared to standard GMAW.Information on power consumption of more traditional metal additive manufacturing techniques, Electron Power consumption compared to metal AM techniques Beam Melting (EBM) of titanium and Direct Metal Laser Sintering (DMLS) of stainless steel, can be found in (Baumers, Dickens, Tuck, & Hague, 2016).

Energy consumption for these printing processes is high: 38.4kWh for EBM and 62.9 kWh for DMLS per kilo of printed material. Energy consumption of WAAM is much lower compared to these techniques. When including 0.75 kW for the exhaust ventilation, the process consumes around 5.18 kWh/kg on average, assuming the print parameters are such that the speed of welding is around half a kg per hour (2 * (1.84 + 0.75) = 5.18 kWh)). The resulting comparison is depicted in Figure 7.

The resulting comparison is depicted in Figure 7. Note that these measurements only include the power required for the shaping processes themselves.) h W k ( g k / n o i t p m u s n o cr e w o P706050403020100) € (m a r g ol i k r e p t s o C1000 900 800 700 600 500 400 300 200 100 0Raw materialWAAMEBMDMLSWAAM, stainless steelEBM, titaniumDMLS, stainless steelFigure 7: Power consumption per kg of metal by WAAM, EBM, and DMLS.Figure 8: Cost estimation per kg of metal by WAAM,EBM, and DMLS.Impacts of the welding gas do not seem to have a large influence on the total (Sproesser et al., 2015). At Shielding gas and emissions the current source of LCIA data, the input materials of GMAW are not given, aside from ‘steel’. Stainless steel creates more environmentally unfriendly, toxic fumes, such as chromium VI, compared to lower alloyed steels.

Stainless steel creates more environmentally unfriendly, toxic fumes, such as chromium VI, compared to lower alloyed steels. However, the significance of these welding emissions is difficult to estimate, as well as the difference with the 413GMAW used in the rough LCA. No measurements have yet been performed due to the required specialized equipment. On the other hand, from (Sproesser et al., 2015) it can be derived that the quantity of emissions is directly correlated to the power consumption. The higher the energy consumption, the more emissions. Concluding, the difference in emissions compared to standard GMAW might have either a positive or a negative impact on the current state of the LCIA.A comparison for metal printing costs is presented in Figure 8. A cost and energy benchmark of EBM and Cost comparison to other AM techniques DMLS can be found in (Baumers et al., 2016).

A cost and energy benchmark of EBM and Cost comparison to other AM techniques DMLS can be found in (Baumers et al., 2016). It highlights the relatively high costs of the Electron Beam Melting of titanium (629 €/kg of which €157 for the raw material) and the Direct Metal Laser Sintering of stainless steel (926 €/kg of which €79 for the raw material), which is largely caused by high machine cost. The cost of WAAM is estimated around 130 €/kg, of which €12 for the raw material (stainless steel), varying with different print parameters. More than half of this share is attributed to labor cost, and can be greatly reduced when the technique is more optimized. Take into account however, that WAAM has a much lower resolution compared to EBM and DMLS, and will need more post-processing.

Take into account however, that WAAM has a much lower resolution compared to EBM and DMLS, and will need more post-processing. Post-processing is not taken into account into this comparison.For further processing, ReCiPe1 was chosen as methodology, a relatively well-known and often-used Selection of methodology and indicator within the LCA domain. For the selection of indicator, the decision was made that the LCA will be expressed in two manners in successive research: a list of (non-normalized) midpoints, plus ReCiPe endpoints. The goal of the list of midpoints is to give more in-depth information to experts. The endpoints are for ease of communication and decision making.

The goal of the list of midpoints is to give more in-depth information to experts. The endpoints are for ease of communication and decision making. Although the uncertainty of endpoint indicators is higher, they provide people without extensive eco-knowledge with an indication.Conclusions and future researchWAAM holds promises for mega-scale constructions, bringing the benefits of digital fabrication to a field where traditional manufacturing techniques limit design and deployment. However, for such a new process, it is important to assess its environmental impact. This is difficult to attain, as aforementioned in the four challenges: • Large differences exist between methodologies and the implied interpretations. • Results can be specified in several ways, posing a trade-off between assessment depth, communication ease,and uncertainty.There is a lack of reliable data of so-called indicators. Even in costly options many significant details are notcovered.

Even in costly options many significant details are notcovered. In order to benchmark an object, at least one comparative object with the same functional unit needs to be found or determined, and assessed in addition to the object itself.Furthermore, we conclude that:In accordance with the ISO 14040 standard (The International Standards Organisation, 2006), the aspect of aesthetics and perceived value will not be taken into account in the assessment in order to keep subjectivity to a minimum.Forerunning results show that the raw material, stainless steel, seems to be the largest contributor to the LCIA. • The power consumption of WAAM is lower compared to average GMAW, and to traditional additive manufacturing techniques per kilo of deposited material.Future research will focus on benchmarking WAAM against alternative manufacturing techniques, including green sand casting and CNC milling.

The goal of this comparison is to give guidelines of what technique is the better choice for the environment in a certain scenario. Firstly, with the aid of the newly acquired SimaPro software and ecoinvent database, a higher quality LCIA of WAAM will be made. Results will be expressed in414(non-normalized) midpoints, and ReCiPe endpoints. Secondly, alternative manufacturing techniques will be assessed, if not available in ecoinvent by researching LCI input and output flows and applying these in SimaPro.AcknowledgementsThis investigation was executed at the 3D Building FieldLab, funded by the Amsterdam institute of Advanced Metropolitan Solutions (AMS). We thank Tim Geurtjens and all engineering staff at MX3D for the support in performing measurements and constructive discussions.ReferencesBaumers, M., Dickens, P., Tuck, C., & Hague, R. (2016). The cost of additive manufacturing: Machine productivity, economies of scale and technology-push.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCHALLENGES IN MAKING COMPLEX METAL LARGE-SCALE PARTS FOR ADDITIVE MANUFACTURING: A CASE STUDY BASED ON THE ADDITIVE MANUFACTURING EXCAVATORAndrzej Nycz*, Mark W Noakes *, Bradley Richardson *, Andrew Messing *, Brian Post *, Jonathan Paul†, Jason Flamm†, Lonnie Love*Oak Ridge National Laboratory, 1 Bethel Valley Rd, Oak Ridge, TN 37831-6479 †Wolf Robotics, 4600 Innovation Drive Fort Collins, CO 80525AbstractThe Additive Manufacturing Excavator (AME) contained several key components that were 3D printed at The Manufacturing Demonstration Facility (MDF) of Oak Ridge National Laboratory (ORNL); it was presented at and performed a live demonstration for the CONEXPO 2017 exhibition in Las Vegas, Nevada in March of 2017.

This paper presents challenges in building functional, large-scale metal parts based on a case study of the excavator. The excavator’s metal arm was 3D printed using a modified Wolf Robotics automated metal inert gas (MIG) welding cell. Tasks included designing a new type of slicer for the metal additive manufacturing (AM) process, integrating the slicing software with the Wolf Robotics system, developing the deposition process, characterizing geometric features and material properties, managing heat, designing mechanical components for metal AM, and developing a machining approach to achieve the final part. Two fully functional excavator arms were printed and machined. Integrated hydraulics passageways that also served as structural stiffeners were included in the build for demonstration purposes.

Two fully functional excavator arms were printed and machined. Integrated hydraulics passageways that also served as structural stiffeners were included in the build for demonstration purposes. As a direct result of this project, Wolf Robotics is now working towards a commercially available large-scale metal AM system.IntroductionThe purpose of this project was to demonstrate the feasibility of implementing large printed metal components on machinery for the construction industry. The main challenges during system development and implementation included deposition process development, geometric feature characterization, robotic path integration and execution, heat management, process appropriate slicing, material properties, mechanical design oriented toward the needs of the AM process, and post-machining to reach necessary interface tolerances.

A substantial development effort was required in a short period of time to ensure success.Arc-based wire feed metal AM was chosen as the best process to produce large metal parts [1]. While metal powder bed printers are available commercially, they are not currently capable of producing large-scale metal parts [2, 3]. Welding-based systems are not as precise as powder bed in terms of print resolution, but there are no inherent limitations to the size of the part that can be printed. Post-process machining of either print technology is typically required for the final product. Therefore, arc-based wire feed technology provided the most cost-effective solution.2024The development timeline was governed by the CONEXPO exhibition beginning in Las Vegas, Nevada on March 7th, 2017. The Wolf Robotics welding equipment was installed at ORNL’s MDF at the end September 2016.

The Wolf Robotics welding equipment was installed at ORNL’s MDF at the end September 2016. During this short amount of time, equipment was setup, safety of the system was assessed, processes were developed, staff members were trained, and the ORNL slicer was adapted to suit the needs of the printing process. A series of smaller parts were printed to establish process parameters such as quality, shape, overlap, adhesion of the weld bead, and print resolution in various Cartesian axes. In order to meet schedule requirements, the printing of the excavator had to begin mid-January 2017 at the latest. During this time, at least two arms needed to be fabricated while leaving enough time for post-process machining, installation, testing, and shipping to Las Vegas.The numerous technical challenges were addressed through extensive literature survey [2, 4].

Significant prior MDF experience in welding and large-scale AM as well as substantial experimental work conducted at the MDF in tungsten inert gas (TIG) and MIG welding-based AM processes was explored. Wolf Robotics and parent company Lincoln Electric provided substantial systems, welding, and materials expertise. The established base of work provided a foundation for successful completion of this project.ChallengesPrint TimeThe expected print time was an important factor throughout the project. Considering the expected weight of the printed arm to be in the 300-400lb range and the maximum rate of deposition of approximately 5lb/h, the expected time of print was in the range of 90-130 hours of continuous deposition. This in conjunction with the timeline affected the design and decision- making process. Any proposed solutions had to address process print time and the available schedule.

This in conjunction with the timeline affected the design and decision- making process. Any proposed solutions had to address process print time and the available schedule. These considerations are similar to what would be required for an industrial implementation of the system [5-7].Deposition Process DevelopmentThe hardware was originally designed for an automated welding process. Therefore, systematic analysis was required to develop a way to additively manufacture large-scale metal parts.One of the first questions to answer is the orientation of the print. The long axis of the part can be vertical (Fig. 1) or horizontal with advantages and disadvantages in both cases. Horizontal orientation significantly decreases the number of layers and causes the bead to become longer. This results in shorter print time and limits the accumulation of the z-direction error.

Horizontal orientation significantly decreases the number of layers and causes the bead to become longer. This results in shorter print time and limits the accumulation of the z-direction error. The downside is, considering the process and print table, decreased temperature of the build, which is far from optimal. It may result in brittleness. The other issue is warping of the build plate causing deformation of the first layers of the part. The vertical orientation increases the print time, but the thermal conditions are closer to optimal. The deformation coming from the base warping was the main factor in deciding to choose the Figure 1. Vertically printedarm on a 4’x3’ table.2025vertical orientation. The z-error accumulation was mitigated by introducing a closed-loop control system and designing extra volume in the part.Metal-arc printing requires adequate print table (Fig 1-2), base plate, and clamps.

Due to the nature of the process, the table and the plate must be made of metal to conduct electricity and be heat resistant. The print table must accommodate the size of the base plate (function of the size of the part) and the size of the mounting solution. In addition, it should have enough thermal capacity to sink and dissipate the energy of the process. It also must be strong enough to withstand the force acting on the build plate due to deformation and warping.The base plate (Fig. 2) must be thick enough to minimize deformation. In this case, the standard thickness was 1 inch. The surface of every plate was thoroughly cleaned to remove traces of oil and contaminants to help with the initial bead deposition and decrease any unwanted electrical resistance. The cleaning also ensured proper flatness of the plate.

The cleaning also ensured proper flatness of the plate. The material used for plates was AISI 1018 mild/low carbon steel.The base plate was mounted to the build plate to preventunwanted motion, ensure electrical contact, and minimize warping during the print (Fig. 2).Gas Metal Arc Welding (GMAW) was used to buildthe excavator stick with Lincoln Electric Power Wave® Technology tuned for low heat input deposition, which allows metal deposits to be stacked. It used a mix of Argon and CO2 at the ratio of 96-4%.A 0.045in diameter (ER70S6) [8] 70ksi low carbon steel wire was used for the feedstock, which provided sufficient deposition fluidity/wetting and mechanical properties. It was also widely commercially available, and the chosen diameter had a good ratio of expected deposition rate and feature resolution. It resulted in a maximum deposition rate of 5.15lb/h for the Surface Tension Transfer (STT) process.

It resulted in a maximum deposition rate of 5.15lb/h for the Surface Tension Transfer (STT) process. For the given process, the arm speed was 6.67mm/s, and the wire feed rate was 84.67mm/s.Figure 2. 3D printed bucket with base plate and mounting clamps. The system was built with a Lincoln Power Wave® S500 welder and AutoDrive® SA servo wire feeder with a secondary push-pull servo motor at the deposition head. This type of wire feeder allows for adequate control of the length of the wire between the surface of the part and the tip of the torch throughout hours of printing.The system was equipped with an arc tracking system, which measured impedance of the welding arc and provided corrections to maintain a specified current. The use of the arc tracking system resulted in consistent contact between the welding tip and the component being fabricated. It also allowed for automated control of the distance of the torch over the part and prevention of part torch collision.

It also allowed for automated control of the distance of the torch over the part and prevention of part torch collision. In addition, another closed loop control system was developed and2026implemented to control the build height of the part. This way the robotic arm was automatically correcting the build height of the part and overriding the command coming from the slicer generated G-Code commands.Geometric Feature CharacterizationIn order to print arbitrary shapes, basic building blocks had to be identified along with their interaction. The process relied on layer and bead structures. Each layer is made up of beads, and beads are considered to be a continuous path of deposition.the following types: insets, skeletons, and infill. The insets are contour lines forming the external and internal edges (Fig. 3). The skeletons are a special type of inset bead that were placed where regular beads were geometrically inadequate.

The insets are contour lines forming the external and internal edges (Fig. 3). The skeletons are a special type of inset bead that were placed where regular beads were geometrically inadequate. The purpose of the infill beads is to fill the empty space between contour beads creating solid metal structure.Beads were dividedintoperimeterFigure 3. Bead types: green – external insets, red internal insets, grey – infill.Unlike polymer deposition, metal AM allows and requires bead-to-bead and layer-to-layer re-melting to build structures of adequate mechanical strength that are free of voids and crack initiation artifacts. A key factor is the choice of proper bead-to- bead overlap and layer heights. The effective distance for contour beads was 4.5mm (center-to-center) resulting in 12mm width for two-bead walls.

A key factor is the choice of proper bead-to- bead overlap and layer heights. The effective distance for contour beads was 4.5mm (center-to-center) resulting in 12mm width for two-bead walls. Based on the arm speed, wire feed rate, nominal process settings, nominal thermal conditions, and two-bead wall structure, the average layer height was 2.31mm.Robotic Path and IntegrationThis system used an ABB 2600 six degree-of-freedom arm with an IRC5 controller (Fig.4). The ABB internal path execution system is inherently incompatible with G-code syntax. Although G-code interpreters do exist [9], the complexity of the task and the level of customization required a new approach.Figure 4. Metal deposition arm.The system automatically performs maintenance activities. They include wire cutting, mechanical nozzle cleaning, nozzle spraying, arm parking, and re-engaging for tip replacement. The wire cut is executed before every bead and after tip change.

They include wire cutting, mechanical nozzle cleaning, nozzle spraying, arm parking, and re-engaging for tip replacement. The wire cut is executed before every bead and after tip change. It guarantees a proper starting process and minimizes voids related to starts and stops. Nozzle cleaning removes the weld spatter and allows for proper shield gas flow. Spraying the nozzle with a liquid agent also decreases the spatter adherence to the torch/nozzle.2027To maintain proper current and tracking calibration, the current conducting tip had to be replaced every 60 minutes. This operation was done manually when the arm automatically placed itself in the maintenance window at predetermined time intervals. The print resumed after the operator restarted the system using the user interface.Most of the process parameters could be controlled on the fly and overwritten as needed by the operator.

The print resumed after the operator restarted the system using the user interface.Most of the process parameters could be controlled on the fly and overwritten as needed by the operator. This allowed for direct experimenting and corrections that are not present on systems executing G-Code exclusively. This included bead tracking and height corrections and could be done both automatically and manually.The path generation system was equipped with a G-Code translation toolbox. It allowed importing G-Code generated geometries and translated them into robot native execution files. In addition, it allowed for offline program simulation and print validationHeat ManagementBased on experience and interviews conducted with welding experts, the goal was to keep the inter-pass temperatures in a range of 200-350C, optimally at about 300C. In the case of a small cross section, a fixed minimum layer time of 240 seconds was enforced to provide enough cooling time.

In the case of a small cross section, a fixed minimum layer time of 240 seconds was enforced to provide enough cooling time. An inter-pass temperature that is too high causes additional temperature rise in external, small areas and in overhang features. This results in melting of the previously deposited layers and may end in catastrophic failure. An inter-pass temperature that is too low affects mechanical properties and may result in brittleness, poor bead- to-bead re-melting, and void creation.The choice of the welding process was also affected by the heat management. A modified Lincoln Electric short arc process was used offering deposition while bead maintaining short arc and low energy. This is accomplished by manipulating the output voltage and current from the welding power supply to control metal droplet size and transfer rate. This resulted in a low heat input process with minimal spatter. It was accomplished through a proper current wave design.excellentFigure 5.

This resulted in a low heat input process with minimal spatter. It was accomplished through a proper current wave design.excellentFigure 5. Path simulation and post processor virtual environmentTo better understand the thermodynamics of large-scale printing using arc welding, IR images (Fig. 5) were taken and temperatures of the part throughout the build history were analyzed. In addition, thermocouples in the based plate and table as well as pyrometers were used to analyze the thermodynamics of printed parts.2028SlicingWhile working on the Big Area Additive Manufacturing (BAAM) development, the MDF team also developed their own slicing software called ORNL Slicer. The Metal BAAM has a similar structure, but not identical, to BAAM. Therefore, a version of the ORNL Slicer meeting the specific needs of metal arc deposition was created. The output of the slicer is G-Code. G-Code is a set of instructions that tells the machine how to build the part.

The output of the slicer is G-Code. G-Code is a set of instructions that tells the machine how to build the part. It includes all the necessary commands for completing the print. This can include temperatures, speeds, moves, and turning devices on and off. (Fig. 6)Figure 6. Path simulation and post processor virtual environmentThe main differences in the metal AM version of ORNL Slicer were the scale of the parts used, complexity, bead definition, start and stop actions, tip wipes, use of bead types, order of print, heating and cooling conditions, curve smoothing tolerance, first layer settings, and direction of printing.Material PropertiesOne of the main concerns of the metal arc technology is the presence of voids and porosity in the build. To test for voids and porosity, several large parts were printed, and their surfaces were machined. The largest part was a 4ft section replica of the excavator arm.

To test for voids and porosity, several large parts were printed, and their surfaces were machined. The largest part was a 4ft section replica of the excavator arm. Over the area of roughly three-square feet, there were no large voids and only a few small (~1mm in diameter) artifacts. Throughout the development process and extensive testing, no cracking was observed. It was credited to the properties of the mild steel ER70S6 wire in conjunction with the chosen process.Figure 7. Sample tensile test resultThe ER70S6 wire is rated at 70 ksi for yield strength[10]. Based on the feedback from the manufacturer, Lincoln Electric, the expected AM yield strength should fall in the 50-60 ksi range. The initial X (printing) and Z (vertical, perpendicular to print) direction tests showed 52 ksi strength (Fig. 7).

The initial X (printing) and Z (vertical, perpendicular to print) direction tests showed 52 ksi strength (Fig. 7). The strength can be affected by heat management or path planning resulting in harder or softer metal.Mechanical DesignLarge-scale metal AM using a MIG process lacks definitive design guidelines in its initial phase. Therefore, based on empirical studies and suggestions from welding experts, in-house rules were defined. The excavator overhang angles were limited to 15 degrees. Single, solid features were limited to one inch in length. Most walls were modeled as two-beads that were 9mm wide with 4.5mmm center-to-center distance. The effective, as printed, distance resulted in 12mm width. This was directed by the necessity of bead overlapping.2029One of the requirements of the project was to print embedded high pressure hydraulic channels within the structure. The standard cross section is circular; however, AM allows for relative freedom in defining shape.

The standard cross section is circular; however, AM allows for relative freedom in defining shape. Therefore, an oval shape was chosen. In addition, the channels were directly integrated into load bearing walls, which minimally increased weight and provided additional stiffness.To account for the transient phase of the process and any initial layer deformations, an extra inch of height was added to bottom of the build. An additional two inches of height were added to the top of the part in case of unexpected height error accumulation.It was decided to post machine the arm and add conventional parts where the benefits of AM would be marginal. The conventional parts included sleeves, bearings, and pins. Those areas were designed as solid surfaces.Figure 8. 4ft test segment arm.A requirement from the machining team was to provide two axes of reference.

Those areas were designed as solid surfaces.Figure 8. 4ft test segment arm.A requirement from the machining team was to provide two axes of reference. The horizontalone was the cut line parallel to the build plate, and the vertical one was a printed feature.In order to test all requirements and assumptions, a four feet section of the excavator arm (inscale 1:1) was printed (Fig. 8).MachiningOne of the advantages of AM is that one can design parts that cannot be produced with traditional machining. On the other hand, one of the disadvantages is that one can design parts with shapes that lack references and are difficult to setup for finish machining operations. Understanding how one is going to machine a part and the needed references is an important step in designing the part.The short test link (Fig. 8) was printed and used as a test article in the workshop.

8) was printed and used as a test article in the workshop. The part was cut off from the base plate, and one side was milled to achieve flat surfaces and to test for voids. Then, the joint areas were cut out, and the fit of pre-machined sleeves was tested. In addition, the test link was used to verify the weldability of the different alloy sleeves to the 3D printed part. No major no susceptibility was observed.issuesandcrackingThe “as printed part” lacked adequate flat surfaces that could be used as lines for CNC machining reference configurations. Therefore, the cut line resulting from removing the base plate was used as a horizontal line for initial setup. The vertical reference line was obtained by manually fitting the part intoFigure 9. Arm on CNC table.2030the digital model CNC path.

The vertical reference line was obtained by manually fitting the part intoFigure 9. Arm on CNC table.2030the digital model CNC path. Proper setup required realigning the parts multiple times because of the lack of adequate references and added time to the machining process.The lessons learned by using the short test link helped to facilitate the machining of the twofull-sized excavator links. Fig. 9 shows one of the links set up on the CNC machine.ResultsTwo identical metal arms were manufactured and post-processed (Fig. 10). One was installed on the excavator and performed a three-day live demonstration during the 2017 CONEXPO Las Vegas show (Fig 10). No mechanical failures were observed. Both the excavator and the extra arm are now located at ORNL’s MDF and are occasionally used for demonstration purposes.Each arm took about five days of nonstop printing; each arm consists of 927 layers and weighs about 400lbs as printed.

Arms were over 7ft tall, and the final height error was close to 1mm. Each arm used approximately 14 miles of wire. The approximated cost of electricity was $20.Figure 10. 3D printed excavator during CONEXPO 2017 demo tests.In addition to the excavator, several other parts were manufactured to demonstrate new AM capabilities. They included a 3ft tall trenching bucket and two 1ft tall bucket teeth.Figure 11. Additively manufactured arm on a deposition table.Since then, a variety of other complex parts have been printed. These include a four-cylinder, full-size engine block, boring disc drill bits for a large-scale hole boring machine, water pump impellers, and complex test geometries.2031The ongoing efforts include metal characterization of printed parts, modeling, and further enhanced control and sensing.

The modeling efforts are focused on thermal, distortion, stress, and metallography prediction conducted with the Simulia (ABACUS) team of Dassault Systems.ConclusionsThis work successfully demonstrated the feasibility of building large-scale metal parts with an arc-based wire feed process. The entire development, analysis, and manufacturing cycle for the excavator arm was completed in four months. Two identical arms were produced. One was mounted on an excavator, and one was mounted in a display stand for inspection and discussion during the trade show. The excavator was successfully demonstrated with nearly three days of continuous digging at CONEXPO.During the initial development process, the need for accurate z-height build control was noted, and a novel closed-loop method for controlling z-height was implemented. The closed-loop method maintained accuracy to 1mm even during large builds.

The closed-loop method maintained accuracy to 1mm even during large builds. Parts did require post machining to remove “support structure” and material near the base layer where weld quality was not ideal. Post machining generally included only those areas necessary required to interface with other, conventional parts in the system. All aspects of development and fabrication were successfully completed. The conclusion of this work was a prototype for a new AM machine and possibly another step toward creation of a main stream economic sector based on large-scale metal AM.References1.2.3. 4.5.6.7.8. 9. 10.Nycz, A., et al., Large Scale Metal Additive Techniques Review. 2016, Oak Ridge National Laboratory (ORNL), Oak Ridge, TN (United States). Manufacturing Demonstration Facility (MDF). Frazier, W.E., Metal additive manufacturing: a review. Journal of Materials Engineering and Performance, 2014. 23(6): p. 1917-1928. Inc., S. Advantages of Wire AM vs. Powder AM.

2015, Springer. p. 39-48. Electric, L., LINCOLN® ER70S-6, T.L.E. COMPANY, Editor. 2016: • Cleveland, OH. Nilsson, D., G-Code to RAPID translator for Robot-Studio. 2016. Company, H.B. Understanding AWS Classifications. 2017; Available from: http://www.hobartbrothers.com/news/268/523/Understanding-AWS-Classifications.html.2032AcknowledgmentsThis manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05- 00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes.

The Department of Energy will provide public access to these results of federally the DOE Public Access Plan research (http://energy.gov/downloads/doe-public-access-plan).sponsoredin accordance withThis material is based upon work supported by the U.S. Department of Energy, Office ofScience, Office of Energy Efficiency & Renewable Energy, Advanced Manufacturing Office, under contract number DE-AC05-00OR22725.2033.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCHANGING PRINT RESOLUTION ON BAAM VIA SELECTABLE NOZZLESPhillip Chesser, Brian K. Post, Randall Lind, Peter Lloyd, and Lonnie J. LoveManufacturing Demonstration Facility, Oak Ridge National Laboratory, Knoxville, TN 37932, USAAbstractBig Area Additive Manufacturing (BAAM) is an additive manufacturing (AM) technique that rapidly deposits polymer to fabricate large components. However, the increase in deposition rates leads to a decrease in resolution and a consequent decline in part surface finish. A novel technique has been developed where the nozzle diameter can be changed mid-print using a poppet nozzle selector. With this technique, a course resolution can be employed to rapidly fabricate the interior of a part, while a fine resolution can be used on the surface.

With this technique, a course resolution can be employed to rapidly fabricate the interior of a part, while a fine resolution can be used on the surface. This allows for improved surface quality and resolution without significantly increasing print time. This work will explain the development of the selectable nozzle and integration with the BAAM system to produce selective high-resolution surfaces on parts.IntroductionFused deposition modeling (FDM) is a common method of additive manufacturing (AM) used to fabricate parts layer by layer from thermoplastic materials. In this AM technique, a filament is heated and softened before being extruded on the part surface. This is a well-established AM method and is good for producing parts with complex and intricate geometries.There are weaknesses with FDM, however. Because filament feedstock is used, the deposition rates are limited. [1] This means that FDM is slow and the build volume is limited.

Because filament feedstock is used, the deposition rates are limited. [1] This means that FDM is slow and the build volume is limited. [2] The largest FDM machines, produced by Stratasys, have a build volume of 3ft x 2ft x 3ft. [3] The throughput and size limitations of FDM have been significantly improved through the development of Big Area Additive Manufacturing (BAAM). [4] BAAM uses pellets as feedstock instead of filament. Pellet feedstock enables deposition rates of up to 100lbs per hour and potentially even higher. The build volume of the largest BAAM is 20ft x 8ft x 6ft.Figure 1 shows a comparison of typical FDM part (left) alongside a BAAM part (right). The FDM part achieves a much higher level of detail and surface finish than the BAAM part.

The FDM part achieves a much higher level of detail and surface finish than the BAAM part. However, the BAAM part was produced in a fraction of the time that the FDM part was produced because of the much higher deposition rates achievable through BAAM.1684Figure 1: FDM and BAAM comparisonThe increase in deposition rates when using BAAM is accompanied by a decrease in print resolution. Typical nozzle diameters for FDM are 0.010in, while typical nozzle diameters for BAAM are 0.3in. This severely limits the level of detail and the surface finish that can be achieved on BAAM parts. In many cases, BAAM parts must be post-possessed to achieve a smooth surface finish. This is commonly done through machining on a 5-axis machining center or router. Machining and post-processing of BAAM parts is typically more time consuming and labor intensive than the initial printing of the part.

Machining and post-processing of BAAM parts is typically more time consuming and labor intensive than the initial printing of the part. The cost involved in the post-processing typically represents the bulk of the cost in parts produced on BAAM.The necessity of extensive post-processing on BAAM parts makes BAAM less attractive for many commercial applications. Therefore, it is desirable to find a solution that increases BAAM part quality and surface finish without significantly affecting deposition rates.One of the most common uses of BAAM is for printing tooling or molds. [5] For example, BAAM has been used to produce molds for composite layups and vacuum assisted resin transfer molding (VARTM). In many of these cases, there are certain surfaces that are critical for the molding operation, while the majority of surfaces are not critical. Figure 2 shows a CAD model of a part produced on BAAM. The model shows a section of the bow for a boat mold.

Figure 2 shows a CAD model of a part produced on BAAM. The model shows a section of the bow for a boat mold. The red represents surfaces critical for molding or for the assembly of the multiple mold components. The total surface area of the part is 24,900 sq in, while the critical surfaces are 3590 sq in. This only represents about 15% of the total surface area of the part. Many of the geometries printed with BAAM have similar ratios of critical surfaces to total surface area. Therefore, it is logical to print the majority of parts at typical BAAM deposition rates, while slowing down deposition rates to print critical areas at a higher resolution.1685Figure 2: Composite layup tool with critical surfaces shown in redMulti-Nozzle Diameter PrintingThe present research focuses on the ability to selectively choose print resolution on the fly.

This enables BAAM technology to print critical surfaces at a fine resolution while printing the rest of the part at a faster, course resolution. Figure 3 shows a diagram of one layer printed with the exterior printed at a fine resolution, and the interior is printed at a course resolution. The orange represents material deposited at a course resolution; the blue represents material deposited at a fine resolution. This enables most the part to be deposited at a course resolution and rapid deposition rates, while the slower, fine resolution deposition is only used on the surfaces of the part where it is needed. In most cases, only some of the part exterior is critical. This results in layers that look more like Figure 4.1686Figure 3: Layer printed at multiple resolutionsFigure 4: Different scheme for multi resolution layerAs finer resolution nozzles are used, the printed part comes closer and closer to the true geometry of the CAD model.

When the CAD part is sliced into layers for printing, surfaces are discretized based on the layer height. This leads to deviation from the specified surface. The larger the discrete layers are, the more error occurs. As the layers are made smaller, the printed part is brought closer to the CAD model geometry. By selectively choosing the nozzle size, the critical surfaces can be printed much closer to the CAD geometry, while non-critical surfaces have larger deviations from the CAD.1687Implementation and TestingTo create a nozzle that can change diameters mid-print, a poppet was added coaxially with the deposition nozzle in the polymer flow near the nozzle tip. Figure 5 shows the poppet (grey) inside of the nozzle (green) in the up position (left) and in the down position (right). In the up position, melted polymer flows around the poppet and out of the course hole in the nozzle’s tip. When the poppet is in the down position, it blocks flow through the larger hole.

In the up position, melted polymer flows around the poppet and out of the course hole in the nozzle’s tip. When the poppet is in the down position, it blocks flow through the larger hole. Melted polymer is now forced through cross drilled holes in the poppet and out a smaller diameter hole in the center of the poppet, which forms the orifice for fine resolution deposition. With this design, a course or fine nozzle resolution can be selected on the fly.While there could be two separate nozzles, there are advantages to being able to select the nozzle diameter using only one nozzle. With a single nozzle, the course and fine resolution deposition is co-located. This simplifies motion control, whereas control would be more difficult if there were two separate deposition heads at different locations. Multiple deposition heads could require multiple extruders as well. This leads to an increase in mechanical complexity.

Multiple deposition heads could require multiple extruders as well. This leads to an increase in mechanical complexity. These problems are resolved by selecting the nozzle diameter within a single nozzle.Figure 5: Nozzle diameter selection poppetAn initial design of a system to implement this is shown in Figure 6. Melted polymer flows from the extruder through a heated block to the nozzle where the selection valve is. Actuation of the poppet was achieved via air cylinders.1688Extruder BarrelAir CylinderNozzleFigure 6: Nozzle offset from the axis of the extruder barrelTo implement the multi-diameter nozzle, printing with a 0.2 inch nozzle was chosen for the course resolution while a 0.1 inch nozzle was chosen for the fine resolution nozzle. It is helpful that the ratio of course-to-fine nozzle diameters be a whole number. This allows for an even number of fine beads to stack up to the height of a course bead.

It is helpful that the ratio of course-to-fine nozzle diameters be a whole number. This allows for an even number of fine beads to stack up to the height of a course bead. Other ratios of course-to-fine nozzle diameters could be tested allowing for a larger discrepancy between the course and fine nozzles.An initial prototype was tested on an existing BAAM platform. Figure 7 shows the prototype setup on the extruder. With this setup, a test part was printed with a 0.2 inch diameter nozzle for the interior and a 0.1 inch diameter nozzle for the exterior. A control part was also printed with the same geometry and a 0.2 inch nozzle diameter for the entire part. Bead heights were half of the bead widths leading to 0.05 inch layer heights with the fine nozzle and 0.1 inch layer heights with the course nozzle.1689Figure 7. First iteration of multi-diameter nozzleThe multi-nozzle part took 3 hours and 48 minutes to print, while the control part took 2 hours and 25 minutes to print.

First iteration of multi-diameter nozzleThe multi-nozzle part took 3 hours and 48 minutes to print, while the control part took 2 hours and 25 minutes to print. This represents a 57% increase in print time for twice the resolution on the part surface. Printing the entire part at the fine resolution would have taken an estimated 9 hours and 5 minutes. Therefore, the multi-resolution print took 58% less time than it would have taken to print the entire part with the high-resolution nozzle. Printing the entire part at a high resolution would not have increased part surface quality at all, over multi-resolution printing. Figure 8 graphically shows the differences in print times and the corresponding resolution of the exterior of the part.1690Print Times100.12980.170.08s r u o H650.06430.0420.02100Course Single ResolutionDual ResolutionFine Single Resolution (Estimated Time)Print TimeExternal Layer HeightFigure 8.

Comparison of Print TimesOn the multi-nozzle part, the entire exterior was printed with a fine resolution. As discussed above, on most parts printed on the BAAM, only a fraction of the surfaces will require the fine resolution. However, the slicing software that prepares the parts for printing does not currently have the capability to selectively print only some surfaces at a fine resolution. Once this ability has been implemented in the software, multi-resolution parts can be printed much faster by only applying the fine resolution to critical surfaces.Photographs comparing the parts are seen in Figures 9 and 10. The fine resolution print visually shows a significant increase in part quality. The layer striation size has been halved due to the doubling of resolution. This leads to a much higher surface quality, as is qualitatively seen in these figures. As research continues, the use of smaller diameter nozzles on the fine resolution setting will be investigated.

This leads to a much higher surface quality, as is qualitatively seen in these figures. As research continues, the use of smaller diameter nozzles on the fine resolution setting will be investigated. This would lead to a further reduction of layer size and consequent further increase in surface quality.Ultimately, it would be ideal to eliminate post-machining for most applications. Instead of machining, critical surfaces could be printed at very fine resolutions, and the minor deviations left in the part could be filled with surface coating techniques. Continued research will work toward achieving this goal, especially through the use of even finer nozzles1691s e h c nIFigure 9: Comparison of multi-nozzle (left) vs. single nozzle (right)Figure 10: Close-up comparison of surface finishesInitial tests reveal some issues that will need to be resolved in future research.

single nozzle (right)Figure 10: Close-up comparison of surface finishesInitial tests reveal some issues that will need to be resolved in future research. First, the starts and stops of extrusions were not clean when printing with the 0.1 inch nozzle because the flow latency of the extruder is amplified on a finer diameter nozzle during starts and stops. This led to significant globing on the seam as shown in Figure 11. Methods of dealing with this issue on the seams have been examined and will be explored in future research.1692Figure 11: Globing on starts and stopsThere are also issues on shallow curves where gaps show the infill underneath the external contours. This is shown in Figure 12. Conventionally, this would be solved by using skins on the top of these sections. The algorithms for adding skins in a part with multiple resolutions is significantly more complex than with single resolution parts.

The algorithms for adding skins in a part with multiple resolutions is significantly more complex than with single resolution parts. While this problem has been explored, these algorithms have not been fully developed and implemented yet.Figure 12: Lack of skins1693ConclusionThe feasibility of selecting print resolution on the fly with a poppet nozzle selector was demonstrated for the BAAM process. It was demonstrated that this method allows an increase in resolution on critical surfaces. A test part was printed with the exterior at twice the resolution of the interior. This allowed for a doubling of exterior resolution without a significant increase in print time over a conventional BAAM part.Future work aims to continue exploring high resolution fabrication while using the lessons learned thus far. Work needs to be done to improve the starts and stops of extrusions. The use of even finer nozzles for the critical sections will also be explored.

Work needs to be done to improve the starts and stops of extrusions. The use of even finer nozzles for the critical sections will also be explored. Software algorithms that allow selection of critical faces to be printed at a fine resolution also need to be developed. Algorithms for improved skins must be developed as well.The potential significance of this work is great. The commercial viability of BAAM for use in producing commercial tooling and molds has been limited because of the expense of post- processing BAAM parts. The ultimate goal of this research is to reduce, or nearly eliminate, the amount of post-processing required.If the resolution of critical surfaces were to come off the printer with a very fine finish, only minor post-processing would be required. This could potentially include minimal manual finishing and light surface coatings. Thus, additively manufactured tooling would be nearly ready to use as it came off the printer.

This could potentially include minimal manual finishing and light surface coatings. Thus, additively manufactured tooling would be nearly ready to use as it came off the printer. This would represent a revolutionary savings of time and money in the tooling industry. The work presented here demonstrates that selection of resolution on the fly is feasible, and it paves the way for moving toward the goal of minimal post-processing of BAAM parts.1694References1. Go, Jamison. "High-Throughput Extrusion-Based Additive Manufacturing" MIT Master of Science Thesis (2015).2. Babu, S., Love, L., Dehoff, R., Peter, W., Watkins, T., & Pannala, S. “Additive Manufacturing of Materials: Opportunities and Challenges.” MRS Bulletin, 40(12), 1154-1161. doi:10.1557/mrs.2015.234. (2015).3. Stratasys. Fortus 900mc. Accessed June 29, 2017. Web. http://www.stratasys.com/3d- printers/production-series/fortus-900mc4. Holshouser, Chris, et al.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCHARACTERISATION OF AUSTENITIC 316LSi STAINLESS STEEL PRODUCED BY WIRE ARC ADDITIVE MANUFACTURING WITH INTERLAYER COOLINGC. R. Cunningham*, J. Wang†, V. Dhokia*, A. Shrokani*, and S. T. Newman*Department of Mechanical Engineering, University of Bath, Bath, BA2 7AY †Department of Architecture and Civil Engineering, University of Bath, Bath, BA2 7AYAbstractWire arc additive manufacturing (WAAM) expands the possibilities of cost effectively producing large-scale, complex metal objects at high deposition rates. Austenitic stainless steel is a commonly used material and has many applications in the marine and nuclear industry due to its high toughness and corrosion resistance.

Austenitic stainless steel is a commonly used material and has many applications in the marine and nuclear industry due to its high toughness and corrosion resistance. Processes parameters such as heat input and interpass temperature can greatly affect the materials properties, part functionality and the economics of WAAM production. However, the effect of these process parameters is not well understood for WAAM of 316LSi. In this research, the effects of the interpass temperature and heat input process parameters on WAAM of austenitic AISI 316LSi stainless steel are experimentally analysed and evaluated. It was found that the heat input and interpass temperature influences the cellular/dendritic morphology and the formation of macro-scale grains within the microstructure.

It was found that the heat input and interpass temperature influences the cellular/dendritic morphology and the formation of macro-scale grains within the microstructure. Additionally, use of higher heat input, resulted in a 28.7% improvement in average Young’s modulus compared to lower heat input, although this remained lower than provided by wrought annealed material.Keywords: Wire arc additive manufacturing, Additive manufacturing, Interpass temperature, 316LSi, Directed energy deposition.IntroductionWire Arc Additive Manufacturing (WAAM) is a directed energy deposition (DED) additive manufacturing approach [1] that uses an electric arc to melt metallic wire feedstock for the production of new parts or repair applications. Compared to other DED technologies that use electron beam or laser beam as a fusion source, the benefits of WAAM include the cost effectiveness of the process.

Compared to other DED technologies that use electron beam or laser beam as a fusion source, the benefits of WAAM include the cost effectiveness of the process. This is due to the ability to source the open architecture equipment required for WAAM from a wide range of suppliers in the arc welding industry. Compared to powder-based additive manufacturing (AM) methods, the wire-based approach in WAAM offers the benefit of cheaper wire feedstock in many materials, and higher deposition efficiency [2]. Notable industrial applications of WAAM in recent years include the first additively manufactured metal bridge [3] and offshore crane hook [4].Austenitic stainless steels find many applications within the marine, construction, nuclear industries due to its excellent corrosion resistance, adequate high temperature mechanical properties, formability and weldability [5].

Although AISI 316L has been intensively investigated as an AM material for powder bed fusion applications [6-11] and has been shown to show unexpectedly high combination of strength and ductility [12], exploration of the material characteristics has been comparatively limited for processing by WAAM.

To date, a number of studies highlighted in table 1 have demonstrated that WAAM can achieve an ultimate tensile strength, yield stress and elongation in excess of cast material and in the region of that provided wrought annealed material across a range of WAAM process types [13-17].426Table 1 Material properties reported for AISI 316L categorised by WAAM process type, specimen orientation and interpass build strategy.WAAM process typeGMAW1GMAW-CMT2 continuousGMAW-CMT pulsedGTAW3-TopTIGGMAW-CMTGMAW – Speed Pulse GMAW – Speed ArcGMAWWrought (annealed) CastSpecimen orientation Vertical Vertical Horizontal Vertical Horizontal Vertical Horizontal Vertical Horizontal Horizontal Horizontal Vertical Horizontal Vertical Horizontal - -Interpass build strategy – dwell duration Fixed dwell – 2 minFixed dwell - unspecified100°C interpass temperature Fixed dwell – 20 s Fixed dwell – 20 s Fixed dwell – 10 s Fixed dwell – 10 s Fixed dwell – 15 s Fixed dwell – 15 s - -UTS (MPa) 533±23 547 577.3 536 588.0 539.9 590.3 537.0 551.0 550.0 ± 6 553.0 ± 2 603.1 557.7 614.4 605.4 480 552YS (MPa) 235±6 336.9 364.3 331.7 374.0 322.2 365.5 313.0 369.0 418.0 417.9 357.6 334.3 361.4 340.7 170 262El (%)48±2 42.0 43.4 45.6 45.1 43.1 42.341%- - - - - 40 551 Gas Metal Arc Welding (GMAW) 2 Cold Metal Transfer (CMT) 3 Gas Tungsten Arc Welding (GTAW)To reliably produce high quality materials and ensure the consistent performance of AM materials in service, the process-structure-property relationship must be well understood [19].

It is also preferable for the desired material properties to be realised in-process due to enhancement of process efficiency. This is particularly beneficial for localized manufacture of large parts, or feature addition or repair applications where sizing and access to a heat treatment furnace might be an issue. To control the material microstructure and properties developed, suitable process parameters must be selected that provide a specific cooling rate, thermal gradient and solidification rate during the build [20].The heat input from the welding torch, based on the current and voltage, directly influences the cooling rate experienced during a WAAM build. However, a range of heat input values are often possible whilst maintaining a stable melt pool and deposition geometry. The advantages of lower heat input in metal AM processes include a refined microstructure, which can translate to improved hardness and tensile strength.

The advantages of lower heat input in metal AM processes include a refined microstructure, which can translate to improved hardness and tensile strength. However, a decrease in residual stress-induced deflection is observed for increasing energy per unit length [21]. High levels of residual stress can have a detrimental effect on fatigue performance and corrosion resistance. Furthermore, higher residual stresses require more robust, and expensive fixturing to constrain the part during the build, the often-cited benefit of tool-less manufacture in additive manufacturing. Distortions that occur during the build, can reduce the quality of the deposition due to variation in contact tip distance. In addition, the volume of δ-ferrite is also affected by the cooling rate in AM. The volume δ-ferrite can be important to control due to its detrimental effect on pitting corrosion resistance [22].

In addition, the volume of δ-ferrite is also affected by the cooling rate in AM. The volume δ-ferrite can be important to control due to its detrimental effect on pitting corrosion resistance [22]. Applications that require low magnetic permeability and high toughness at cryogenic temperatures may also require δ-ferrite to be controlled to acceptable levels through WAAM process.The strategy employed for build temperature management can also affect the cooling rate and part quality. A fixed or time-variant dwell period may be used to set the inter-layer cooling time and prevent undesired heat accumulation that can destabilise the deposition geometry. Alternatively, an active cooling process may be used to increase the cooling rate post-deposition [23].

Alternatively, an active cooling process may be used to increase the cooling rate post-deposition [23]. By increasing the interpass temperature a significant reduction in baseplate distortion was found for laser DED of Ti-6Al-4V control due to greater recrystallisation with associated dynamic reduction427Ref[13][14][15][16][17][18]of plastic strain and also stress relaxation at the higher build temperature [24]. As shown in table 1, in all cases a fixed or time variant interpass temperature dwell was used, indicating that this is an important aspect of WAAM 316L/LSi production. However, except for Wu et al.

As shown in table 1, in all cases a fixed or time variant interpass temperature dwell was used, indicating that this is an important aspect of WAAM 316L/LSi production. However, except for Wu et al. (2019) [17] where the effect of a small difference in dwell periods was considered, the effect of interpass dwell on the structure-property characteristics of the deposited material has not been explored.The aim of this research study is to investigate the effect of heat input and interpass temperature on the WAAM location-specific material properties in relation to conventionally manufactured material.Material fabricationThin-wall samples of length 150 mm and approximately 100 mm height as shown in Fig. 1 were produced in this study using a parallel path deposition strategy. Wire of type and 1 mm diameter was deposited onto 316L substrates of 50 mm × 200 mm × 6 mm dimension. The chemical composition of the wire and substrate is shown in table 2.

Wire of type and 1 mm diameter was deposited onto 316L substrates of 50 mm × 200 mm × 6 mm dimension. The chemical composition of the wire and substrate is shown in table 2. Prior to commencing each build, the substrate surfaces were polished and then cleaned with acetone.

The substrate was fixed during the WAAM build with bolts in each corner of the substrate.Figure 1 As-built geometry of 316LSi thin-walls and co-ordinate system.Table 2 Chemical compositions of the wire and substrate used in the wire arc additive manufacture of the test samples.CrNiMoCMnSSiNCuWire - 316LSi18.25012.0702.5300.0081.6000.01100.7900.03200.083Substrate - 316L17.01010.1072.1520.0231.2320.00490.4600.0507The factors of interpass temperature and heat input were investigated in a full factorial experimental design shown in table 3 with test walls built from process parameter combinations of low interpass temperature with low heat input (LIT-HHI), low interpass temperature with low heat input (LIT-LHI), high interpass temperature with high heat input (HIT-HHI), and high interpass temperature with low heat input (HIT-LHI).The heat input levels to be investigated were derived from the minimum and maximum travel speeds that could be achieved without excessive build-up of the weld bead or humping respectively to prevent any instabilities in deposition geometry during deposition and were achieved by varying the travel speed of the weld torch only.

The upper level of interpass temperature was determined by the maximum temperature possible to achieve passively with continuous deposition for the lower heat input condition.428FeBal.Bal.Table 3 Experimental wire arc additive manufacturing build parameters.Experiment Ref.

LIT-HHI LIT-LHI HIT-HHI HIT-LHIInterpass temperature (°C) Heat Input (kJ/mm) Travel speed (mm/s)50 (LIT)160 (HIT)0.47 (HHI) 0.26 (LHI) 0.47 (HHI) 0.26 (LHI)2.32 4.90 2.32 4.90The build parameters were fixed throughout all of the experiments and the values are provided in table 4.Table 4 Fixed wire arc additive manufacturing build parametersParameters Wire feed rate Contact tip distance Shielding gas flow rate Shielding gas composition Welding angle Wire diameter Pulse type Pulse frequencyValues 2 m/min 7.5 mm 20 L/min 97.5%Ar/2.5%CO2 0° 1 mm Primary 5 HzThe WAAM equipment consisted of an Aristo 4004i power supply and wire feeder (ESAB Group (UK) Ltd), a fume extractor, controller and shielding gas supply configured as shown in Fig. 2. The weld torch cable which supplies the wire and shielding gas was routed through the WAAM enclosure, where it was affixed to the final motion system stage along with a pyrometer (Micro- Epsilon, CTM-2CF75-C3).

The pyrometer provided continuous feedback of the surface temperature 30 mm in front of weld torch in the travel direction and was driven to the weld start point between passes to monitor the interpass temperature. The temperature readings were calibrated with a k-type thermocouple, and accuracy of the readings was monitored at regular intervals throughout the experimental builds.HardwareWAAM enclosure 3 Axis motion system Pyrometer to monitor build temperatureFume extractorSoftware instructions Motion Arc ON/OFFIIIaSoftware Monitoring lnterpass Temperature. -Figure 2 WAAM experimental set up.429Materials AnalysisTo examine the microstructure through the build height of the experimental thin-walls, a section of 4 mm thickness was extracted from the location shown in Fig. 3 using a conventional band-saw. These sections were polished mechanically with SiC from grit number 400 to 4000, and subsequently polished with 0.5μm colloidal silica.

3 using a conventional band-saw. These sections were polished mechanically with SiC from grit number 400 to 4000, and subsequently polished with 0.5μm colloidal silica. A chemical etch using Marble’s reagent solution (4g CuSO4, 20ml HCl, and 20ml H20) was performed for 10 s to provide contrast between the γ- austenite and δ-ferrite phases. The YZ surface of each specimen was examined using an optical microscope and scanning electron microscope (SEM) (JEOL JSM-6480LV) was used to analyse the fracture surfaces of the tensile specimens. To investigate the variation in phase content and morphology in each sample, images were made at a magnification x 1000 with a spatial scanning frequency of 1 mm along the build height in YZ.

To investigate the variation in phase content and morphology in each sample, images were made at a magnification x 1000 with a spatial scanning frequency of 1 mm along the build height in YZ. These images were converted to binary using ImageJ and the ratio of white to black pixels recorded representative of the γ-austenite to δ-ferrite content respectively22,575 30Units (mm)IH3' 'V1H2V2 V3ISectio n for micros! ructural analysis and mi testingera-hardness,; ,H1~ ~, ;iFigure 3 Orientation of tensile specimens within the WAAM experimental walls and specimen dimensionsThe tensile coupon specimens were water jet cut from the experimental thin-walls in the orientations as shown in Fig. 3. The specimen dimensions were specified in accordance to ASTM E8 [25], with all specimens CNC machined to a uniform thickness of 3.3 mm.

3. The specimen dimensions were specified in accordance to ASTM E8 [25], with all specimens CNC machined to a uniform thickness of 3.3 mm. Uniaxial tensile tests were performed at room temperature using an Instron 3369 50 kN load cell with a clip gauge extensometer to study the plastic deformation behaviour on the macroscopic scale. The extension strain rate used was 0.00007 s-1 in the elastic deformation region, followed by 0.00024 s-1 upon transition to plastic deformation.The experimental data was processed to determine the Young’s modulus, yield strength, the ultimate tensile strength and fracture elongation. The Young’s modulus was determined by considering the linear part of the stress–strain curve, and the yield strength adopted the 0.2% proof stress. Statistical analyses were carried out to identify significant effects and interactions via three- way analysis of variance (ANOVA) tests using Minitab® 18 software.

Statistical analyses were carried out to identify significant effects and interactions via three- way analysis of variance (ANOVA) tests using Minitab® 18 software. The factors considered were the interpass temperature, heat input (quantitative factors) and specimen orientation (categorical factor). Main effects and interactions were determined to be significant if p-value was less than 0.05. The data was also checked for violations of normality and equivalence of variance as required for performing ANOVA.The variation in micro-hardness throughout the build height was established with Vickers micro- hardness test equipment (Leco M400) with the same samples that were used to examine the430microstructural properties In accordance the ISO 6507-1 [26], a load of 500 g was applied for 10 s, with the first indentation located 3 mm inwards from the bottom surface of substrate and at the centreline of the WAAM layers.

Subsequent measurements were made along the sample in increments of 1 mm in the build height direction.The distortion of the substrates following release from the build fixture was captured using a ROMER Absolute scanning arm (Hexagon Manufacturing Intelligence, UK) and Geomagics Wrap(cid:147) software (3D Systems Inc). The data was processed to obtain the profile of maximum deformation in the XZ and YZ planes.Microstructure and periodicityThe LIT-LHI and HIT-HHI thin-wall specimens showed large irregularly shaped columnar macro- grains which grew across the layer bands in the build direction (BD) as shown in Fig 4. The LIT- LHI columnar grains displayed the greatest aspect ratio due to the strong thermal gradient developed by the combination greater inter-layer cooling and low heat input deposition.

The LIT- LHI columnar grains displayed the greatest aspect ratio due to the strong thermal gradient developed by the combination greater inter-layer cooling and low heat input deposition. The reduced melt-pool temperature also minimizes the time for heterogeneous nucleation to occur ahead of the solidification front, encouraging epitaxial growth dominant along the <100> easy growth direction due to the strong thermal gradient. The HIT-HHI thin-wall specimen, experienced the most elevated build temperatures and consequently had a strong tendency for grain growth [27].

The HIT-HHI thin-wall specimen, experienced the most elevated build temperatures and consequently had a strong tendency for grain growth [27]. However, with less directional heat transfer along the build height, more diverse competitive grain growth was able to occur, reducing the length and aspect ratio of the grains that formed to impart a less extreme texture.a)BDb)BDFigure 4 Columnar grain growth visible through the macro-structure in a) LIT-LHI and b) HIT-HHI experimental thin walls.The as-built cross section morphologies of the banding zones at 15 mm from the bottom of the substrate are shown in Fig. 5, respectively for each of the experimental thin-walls. It was evident from each of the experiments that the δ-ferrite developed was of dendritic morphology within an austenite matrix. Three periodic zones are visible, including a fine grain zone (FZ) immediately to431the left of the re-melt banding line caused by recrystallisation.

Three periodic zones are visible, including a fine grain zone (FZ) immediately to431the left of the re-melt banding line caused by recrystallisation. This was primarily lathy δ-ferrite morphology, although for HIT-HHI and LIT-LHI cellular regions were also observed. The transition zone (TZ) displayed a progressive coarsening of this FZ microstructure due to increasing re-heat effect from the superimposed layers in this direction and conversely a progressive refinement of the coarse grain zone (CZ) due to diminishing re-heat effect. The CZ was positioned to the right of the re-melt banding line, received the most re-heat without re-melting and was subsequently of the most coarsened vermicular morphology.Dendritic δ-ferriteDendritic δ-ferriteDendritic δ-ferriteDendritic δ-ferriteFigure 5 Experimental thin wall morphologies at the banding zone 15mm from the bottom of the substrate for a) LIT-HHI b) LIT- LHI, c) HIT-HHI d) HIT-LHI.

Note: Vickers Micro-hardness indentations are visible in c) and d) are circled in black.The δ-ferrite network within LIT-HHI (Fig 5a inset) was observed to be more extensive with larger primary dendrite arm length, whereas the δ-ferrite is discontinuous in CZ of the HIT-LHI (Fig 5d inset). This indicated that greater decomposition of δ-ferrite occurred, possibly due to greater reheating effect due to reduced layer height or increased interpass temperature. It is known that432vermicular δ-ferrite morphology transforms to fine short ferrite precipitates when subject to repeated repair welding of 316L [28]. The similarity of the process to WAAM and the lengthy re- heat distance for HIT-HHI can explain the relatively fine microstructure observed within the macroscale grains. As the morphology within the FZ of LIT-LHI could not be observed clearly with optical microscopy, this sample was analysed with SEM. As shown in Fig.

As the morphology within the FZ of LIT-LHI could not be observed clearly with optical microscopy, this sample was analysed with SEM. As shown in Fig. 6, colonies of cells were present which were grouped into sub-grains, with a sub-grain boundary highlighted. Lathy dendritic primary arm δ-ferrite could be observed within the fine cellular austenitic matrix.Figure 6 SEM (BEI) micrograph of 85mm build height of LIT-LHI showing the lathy δ-ferrite within cellular austenite and a sub- grain boundary.The LIT-LHI and HIT-HII demonstrated the most consistent average white pixel concentration along the built height as shown in Fig. 7, which was attributed to the more refined and homogeneous morphology between the CG, TZ and FZ for these samples. Overall, the LIT-LHI displayed the greatest volume of δ-ferrite which was expected due to the high cooling rate achieved with the combination of low heat input and interpass temperature.

Overall, the LIT-LHI displayed the greatest volume of δ-ferrite which was expected due to the high cooling rate achieved with the combination of low heat input and interpass temperature. These conditions would lead to the most retained metastable δ-ferrite and least opportunity for decomposition through the reheat cycles.LIT-HHIHIT-HHILIT-LHIHIT-LHI0.9n o i t a r t n e c n o c0.80.70.6le x pie0.50.4t ih w e g a r e v A0.30.20.10010203050 Wall height (mm)4060708090100Figure 7 Average white pixel concentration along the WAAM experimental wall height.

Etching and the converting to binary process resulted in austenite = white, δ-ferrite = black.433Mechanical properties and deformation behaviorThe tensile properties of AISI 316LSi samples fabricated by WAAM were evaluated as shown in table 5 with the data are expressed as mean ± standard deviation from sample group number, n.Table 5 Young’s modulus, yield strength, ultimate tensile strength, and elongation of the investigated as-built materials including standard deviation within sample group of size, n.Experiment ref.

The horizontal specimens provided an average elongation of 45.2% compared to 40.0% provided by the vertical specimens. This indicates that on average the specimens extracted in the horizontal orientation provide greater ductility. Previous studies investigating ductility of 316L have shown that when more grain boundaries are encountered, higher strength and lower elongation are expected [29]. A higher number of grain boundaries and lower elongation are expected in horizontal specimens due to directional grain growth, suggesting that another mechanism is providing the additional ductility in WAAM horizontal specimens. The heat input was found to have a significant effect on the Young’s modulus, with high heat input providing an average of 164.3GPa, compared to low heat input with an average of 127.7GPa. This could be significant for high strength applications as in both cases plastic deformation would proceed at a lower load than for wrought annealed material.

This could be significant for high strength applications as in both cases plastic deformation would proceed at a lower load than for wrought annealed material. Besides Young’s modulus and elongation, the other mechanical properties investigated generally exceed that provided by wrought annealed material.As identified in other research by Yang et al. [30], there is a general decrease in micro-hardness with increasing distance from the substrate. The high heat input level with high interpass temperature substantially reduced the micro-hardness for the start section of build height compared to the same heat input with longer cooling between passes. In contrast, for the low heat input samples there was little difference in start section micro-hardness for both interpass temperatures. This is reflected in the similarity of the microstructure as shown in Fig. 8 and is explained by the limited heat accumulation in the high interpass temperature at the start of the build.

This is reflected in the similarity of the microstructure as shown in Fig. 8 and is explained by the limited heat accumulation in the high interpass temperature at the start of the build. As the build proceeds, the effect of the high interpass temperature becomes more apparent with the low heat input experiments, with the high interpass temperature negatively impacting on the micro-hardness within the top section of the build height.

9. The specimens all show typical fracture features of ductile materials, with small dimples of less than 1 μm in diameter visible throughout the surface.LIT-LHIHIT-LHILIT-HHIHIT-HHI3 H1 HFigure 9 SEM micrographs of the surface of the fractured tensile specimens435Whilst no significant differences in the morphology of dimples could be observed between the experimental test conditions, the low interpass temperature with the high heat input condition (LIT- HHI) appeared to reduce the dimple diameter compared to the high interpass temperature condition. Future work will explore possible correlations of dimple size and frequency, with the tensile properties of 316LSi produced by WAAM for each of experimental test conditions investigated in this research.DistortionThe profile of maximum distortion in the XZ and YZ planes for each wall are shown in Fig. 10.

10. It can be seen that in the XZ plane, a high interpass temperature can reduce the residual stress induced distortion by a small amount, with the greatest improvement seen for high heat input experiments. Reduction in distortion was also found with increased interlayer cooling time for the lower heat input experiments in the YZ plane. However, the results are less conclusive for high heat input condition where the high interpass temperature distortion (HIT-HHI) exceeds that of the low interpass temperature (LIT-HHI) on one side of the YZ plane.

This indicates that a high interpass temperature build strategy may not be as effective at reducing distortion for the YZ plane with higher levels of heat input.a)2.0b) 2 .0E 1.5 -S ~ .5 1.0 C 0 t ~ i5 05E1 .5 -S N >< .5 1.0 C 0 :e ~ i5 05II'II0.00.00100 Distance along substrate (mm)501502000100 Distance along substrate (mm)50150c)1.0d) 1.0~0 .8 E -S ~ 0.6 .5 C:e 0.42 VJ i5 0.2' ' ' '' ' ' ' ' ' ' ' '' '' '~0 .8 E -S ~ 06 .5 C:e 0.42 VJ i5 0.2----.,..,..,..,..,..,.0.020 20 0 Distance from deposition centre-line (mm)10100.020 20 0 Distance from deposition centre-line (mm)1010Figure 10 Maximum distortion profiles in the XZ plane for a) HIT-HHI and LIT-HHI b) HIT-LHI and LIT-LHI and in YZ plane for c) HIT-HHI and LIT-HHI d) HIT-LHI and LIT-LHI.436200ConclusionsThis paper has experimentally investigated the effect of heat input and interpass temperature on the microstructure, hardness, tensile strength, and stiffness for stainless steel 316LSi walls manufactured by WAAM identifying the following new findings and conclusions:1.

The microstructures of AISI 316LSi produced by WAAM generally consisted of a mixture of fine columnar dendrites and isolated cellular regions. The microstructures varied periodically across the layer bands, with fine, transitional and coarse grain regions.2. The inter-layer cooling time between layers affected the development of macro-scale grain structures within the microstructure. These exceeded 5 mm with the low interpass temperature and low heat input condition. Macro-scale grains were also observed with the high interpass temperature with high heat input experiment however, these were more equiaxed which was attributed to break-down by competitive grain growth.3. The heat input was found to have a significant effect on the Young’s modulus of AISI 316LSi produced by WAAM with high heat input processing parameter providing an average of 164.3GPa, compared to low heat input with an average of 127.7GPa.4.

The tensile specimens extracted in the horizontal orientation provide statistically significant greater ductility, with an average elongation of 45.2% compared to 40.0% provided by the vertical specimens. Inter-layer cooling time influences the distortion of WAAM produced parts, with results indicating that a higher interpass temperature could be used to reduce the residual stress induced distortion of the substrate, although further research is required. AcknowledgementsThe authors would like to acknowledge support from Renishaw PLC (EPSRC CASE studentship No. 1780168).References[1] [2][3] [4][5][6][7][8]ASTM F3187-16, Standard Guide for Directed Energy Deposition of Metals, 2016. A. Nycz, A. I Adediran, M. W. Noakes, and L. J. Love, "Large Scale Metal Additive Techniques Review," Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference, pp. 2001-2006, 2016. MX3D.

S. A. Khairallah, A. T. Anderson, A. Rubenchik, and W. E. King, "Laser powder-bed fusion additive manufacturing: Physics of complex melt flow and formation mechanisms of pores, spatter, and denudation zones," Acta Materialia, 108, 36-45, 2016. Y. Zhong, L. Rannar, L. Liu, A. Koptyug, S. Wilman, J. Olsen, D. Cui, and Z. Shen, "Additive manufacturing of 316L stainless steel by electron beam melting for nuclear fusion applications," Journal of Nuclear Materials, 486:234-245, 2017.[12] M. S. Pham, B. Dovgyy, and P. A. Hooper, "Twinning induced plasticity in austenitic stainless steel 316L made by additive manufacturing," Materials Science and Engineering: A, 704: 102-111, 2017. X. H. Chen, J. Li, X. Cheng, B. He, H. M. Wang, and Z.

X. H. Chen, J. Li, X. Cheng, B. He, H. M. Wang, and Z. Huang, "Microstructure and mechanical properties of the austenitic stainless steel 316L fabricated by gas metal arc additive manufacturing," (in English), Materials Science and Engineering a-Structural Materials Properties Microstructure and Processing, 703: 567-577, Aug 4 2017. N. Rodriguez, L. Vázquez, I. Huarte, E. Arruti, I. Tabernero, and P. Alvarez, "Wire and arc additive manufacturing: a comparison between CMT and TopTIG processes applied to stainless steel," Welding in the World, 62(5): 1083-1096, 2018. A. Queguineur, G. Rückert, F. Cortial, and J. Y. Hascoët, "Evaluation of wire arc additive manufacturing for large-sized components in naval applications," Welding in the World, 62(2):259- 266, 2017. L. Wang, J. Xue, and Q.

Cortial, and J. Y. Hascoët, "Evaluation of wire arc additive manufacturing for large-sized components in naval applications," Welding in the World, 62(2):259- 266, 2017. L. Wang, J. Xue, and Q. Wang, "Correlation between arc mode, microstructure, and mechanical properties during wire arc additive manufacturing of 316L stainless steel," Materials Science and Engineering: A, 751:183-190, 2019.[13][14][15][16][17] W. Wu, J. Xue, L. Wang, Z. Zhang, Y. Hu, and C. Dong, "Forming Process, Microstructure, and Mechanical Properties of Thin-Walled 316L Stainless Steel Using Speed-Cold-Welding Additive Manufacturing," Metals, 9(1): 109, 2019. ASM International, Metals Handbook. 9th Ed. Vol. 3. Properties and Selection Stainless Steels, Tool Materials & Special-purpose Metal. ASM, 1980. T. DebRoy, H. L. Wei, J. S. Zuback, T. Mukherjee, J. W. Elmer, J. O. Milewski, A. M. Beese, A. Wilson-Heid, A. De, and W.

ASM, 1980. T. DebRoy, H. L. Wei, J. S. Zuback, T. Mukherjee, J. W. Elmer, J. O. Milewski, A. M. Beese, A. Wilson-Heid, A. De, and W. Zhang, "Additive manufacturing of metallic components – Process, structure and properties," Progress in Materials Science, 92: 112-224, 2018.[18][19][20] W. J. Sames, F. A. List, S. Pannala, R. R. Dehoff, and S. S. Babu, "The metallurgy and processing science of metal additive manufacturing," International Materials Reviews, 61(5): 315-360, 2016. A. S. Wu, D. W. Brown, M. Kumar, G. F. Gallegos, and W. E. King, "An Experimental Investigation into Additive Manufacturing-Induced Residual Stresses in 316L Stainless Steel," Metallurgical and Materials Transactions A, 45(13): 6260-6270, 2014. C. D. Lundin, W. T. Delong, and D. F. Spond, "Ferrite-Fissuring Relationship in Austenitic Stainless-Steel Weld Metals," (in English), Welding Journal, 54(8): S241-S246, 1975. P. Henckell, K. Günther, Y. Ali, J. P. Bergmann, J. Scholz, and P.

P. Henckell, K. Günther, Y. Ali, J. P. Bergmann, J. Scholz, and P. Forêt, "The Influence of Gas Cooling in Context of Wire Arc Additive Manufacturing—A Novel Strategy of Affecting Grain Structure and Size," in TMS 2017 146th Annual Meeting & Exhibition Supplemental Proceedings, M. The Minerals and S. Materials, Eds. Cham: Springer International Publishing, 2017, 147-156. E. R. Denlinger, J. C. Heigel, P. Michaleris, and T. A. Palmer, "Effect of inter-layer dwell time on distortion and residual stress in additive manufacturing of titanium and nickel alloys," (in English), Journal of Materials Processing Technology, 215: 123-131, 2015. ASTM E8/E8M, Standard Test Methods for Tension Testing of Metallic Materials, 2013. BS EN ISO 6507-1, Metallic materials – Vickers hardness test Part 1: Test method (ISO 6507- 1:2018). K. Saeidi and F. Akhtar, "Subgrain-controlled grain growth in the laser-melted 316 L promoting strength at high temperatures," R Soc Open Sci, 5(5): 172394, 2018. I.

K. Saeidi and F. Akhtar, "Subgrain-controlled grain growth in the laser-melted 316 L promoting strength at high temperatures," R Soc Open Sci, 5(5): 172394, 2018. I. AghaAli, M. Farzam, M. A. Golozar, and I. Danaee, "The effect of repeated repair welding on mechanical and corrosion properties of stainless steel 316L," Materials & Design (1980-2015), 54: 331-341, 2014.[21][22][23][24][25] [26][27][28]438[29][30]Y. Zhong, "Sub-Grain Structure in additive manufactured SS316L," PhD thesis, Department of Materials and Environmental Chemisty, Stockholm University, 2017. N. Yang, J. Yee, B. Zheng, K. Gaiser, T. Reynolds, L. Clemon, W. Y. Lu, J. M. Schoenung, and E. J. Lavernia, "Process-Structure-Property Relationships for 316L Stainless Steel Fabricated by Additive Manufacturing and Its Implication for Component Engineering," Journal of Thermal Spray Technology, 26(4): 610-626, 2016.439.

ARACTERISATION OF METAL DEPOSITION DURING ADDITIVE MANUFACTURING OF Ti-6Al-4V BY ARC-WIRE METHODSN. P. Hoye*, E. C. Appel*, D. Cuiuri*,# and H. Li*,#School of Mechanical, Materials and Mechatronic Engineering, University of Wollongong, NSW, Australia 2518 #Defence Materials Technology Centre, University of Wollongong, NSW, Australia 2518AbstractThis study considers the use of the gas tungsten arc (GTA) welding process in conjunction with ‘cold’ wire addition to give layer-wise build-up of thin walled structures, simulating those commonly found in aerospace applications, which may in the future be manufactured by additive means. Taguchi DOE and multiple regression analysis methods have been applied to quantitatively establish relationships between common process parameters including arc length, arc current, travel speed and wire feed speed and resulting weld bead geometries and actual metal deposition rates.

Mathematical expressions for build-up height, thickness and surface roughness are presented and evaluated against experimental data, with observations related to physical phenomena.IntroductionSince their initial introduction and development, additive manufacture or solid freeform fabrication techniques have seen an ever increasing degree of interest in the field of engineering manufacturing. In general, this group of processes offers highly flexible, cost effective alternatives to conventional manufacturing methods for both large production runs and one-off prototype creations. There is particular interest for the adoption of such techniques into the aerospace sector where the increased material utilisation has the potential to deliver significant financial benefits when using costly materials such as titanium alloys [1, 2].

One such method of additive manufacture, considered in the present study, is gas tungsten arc (GTA) welding with mechanised ‘cold’ wire feed where weld beads are laid successively onto a substrate giving layer-wise build-up of the desired profile.It is widely known that the parameters employed during a welding process have a great influence over the geometry of the resulting weld beads, with these geometries in turn influencing the properties of the welded structure. The importance of these weld bead geometries may be considered to be especially important during additive manufacturing procedures, such as the one considered in this study, where each bead combines to yield a final near net-shape freeform with a predefined desired profile. Any significant deviation from this desired profile may be considered as inefficiency in the process, requiring additional processing such as increased material removal and wastage or re-working, which would contribute unnecessary costs.

It is clear then that a thorough understanding of the relation between process parameters and resulting weld bead geometries is crucial for the optimisation weld-based additive manufacturing procedures.1015Often times welding procedures are developed by a method of trial and error where process parameters are varied until an acceptable outcome, in terms of weld bead geometry and properties, is achieved. While such methods can deliver acceptable results, they are time- consuming and rarely yield an optimised solution. In contrast, the development of mathematical models relating input parameters to output variables allows for process optimisation across a wide range of scenarios as well facilitating process automation to further increase efficiency [3, 4].

Much literature exists describing models for various aspects of the GTA welding process including weld pool and weld bead geometries [5-9], however these studies are primarily concerned with conventional joint configurations such as butt and fillet welds and are not specific to titanium alloys. As such, these models may not be well suited for predicting the geometries of multi-layer weld deposits used for additive manufacture of components from titanium alloys.

As such, these models may not be well suited for predicting the geometries of multi-layer weld deposits used for additive manufacture of components from titanium alloys. For the specific case of additive manufacture with titanium alloys, the work of Charles [10] presents thermal models for prediction of microstructure evolution using GTA welding, while Sequeira Almeida and Williams [11] discuss the development of process models for modified GTA and gas metal arc (GMA) welding processes.The aim of the present study was to establish a comprehensive understanding of the relationship between common process parameters and resulting weld deposit geometries in the additive manufacture of titanium alloys using conventional GTA welding. This was done initially through experimentation with the effect of various input parameters on deposit geometry observed and quantified.

This was done initially through experimentation with the effect of various input parameters on deposit geometry observed and quantified. Using this data a series of mathematical models for the subsequent build-up profile were generated through analysis of variance (ANOVA) and multiple regression techniques. It is intended that such models will allow for the optimisation of process parameters during the additive manufacture of titanium alloys by GTA welding and so validate the technique as a possible low cost alternative to more conventional manufacturing methods used in the aerospace sector.ExperimentalSamples of commercially sourced Ti-6Al-4V plate with dimensions 100mm x 16mm x 9.65mm were used to form the substrate onto which weld beads were deposited using Ti-6Al-4V wire with a diameter of 1.0mm. The nominal chemical composition of these materials (in wt%) is provided in Table 1.Table 1.

The nominal chemical composition of these materials (in wt%) is provided in Table 1.Table 1. Chemical composition of Ti-6Al-4V alloy.AlloyTiCFeComposition (wt%)HNOAlVTi-6Al-4V99.20.10.30.0150.030.256.14.0Welding was conducted using water cooled Conley & Kleppen (CK) machine mount torch coupled to a Kemppi MasterTIG MLS 2000 inverter power supply with independent wire feed provided through a CK WF3 dedicated cold wire feed unit. Inert gas shielding was achieved with welding grade pure argon using pre and post flow options in addition to a custom fabricated trailing shield. Shielding gas flow rates and other common process parameters are detailed in Table 2.1016Table 2.

Shielding gas flow rates and other common process parameters are detailed in Table 2.1016Table 2. Summary of GTA welding parameters.Polarity Electrode Shielding gas Flow rate – torch nozzle Flow rate – trailing shield front Flow rate – trailing shield rear Pre-flow duration Up slope duration Down slope duration Post flow durationDCEN 2% Ceriated, 2.4mm Ø Welding Grade Argon 8 L/min 10 L/min 7 L/min 3 seconds 2 seconds 1 second 30 secondsFor the present study, five common input parameters of the GTA welding process, namely arc length (L), arc current (I), travel speed (TS), wire feed speed ratio (WFSR) and interpass temperature (IT), were selected to be varied in order to characterise the geometry of deposited weld beads. Each parameter was examined at four different levels, selected based on previous experimental procedures, in order to cover a broad operating envelope. This gives 3 degrees of freedom (DOF) per parameter; totalling 15 DOF’s for the five process parameters.

This gives 3 degrees of freedom (DOF) per parameter; totalling 15 DOF’s for the five process parameters. The levels for each parameter are given in Table 3. Using the design of experiment (DOE) methods of Taguchi, a total of 16 experiments were conducted with the L’16 orthogonal array used to determine the combination of parameters for each experiment.Table 3.

Definition of input process parameters and levels used for GTA welding experiments.Parameter (units) Level 1 Level 2 Level 3 Level 4Arc Length (mm) 3.0 3.5 4.0 5.0Arc Current (amperes) 100 120 140 80Wire Feed (ratio to TS*) 5 7 9 11Travel Speed (mm/min) 100 150 200 250Interpass Temp (°C) 40 90 150 200Wire feed was defined by a ratio against travel speed in order to eliminate impractical combinations of the two.Samples were held ‘on edge’ in a steel clamping arrangement to simulate the beginning of a wall-like build-up, with the entire fixture attached to a linear actuator, above which the torch and wire feed were held stationary. Experiments were performed with ten successive passes conducted for each parameter set to yield a small vertical wall. This geometry was selected to mimic the type of build-up that would be required to manufacture thin walled, pocketed sections commonly found in structural aerospace applications.

This geometry was selected to mimic the type of build-up that would be required to manufacture thin walled, pocketed sections commonly found in structural aerospace applications. Arc current and voltage (V) were observed for each pass using a Triton Electronics AMV4000 weld monitor, allowing the determination of arc energies (AE) and heat inputs (HI). Temperature profiles for each pass were recorded through NI LabVIEW software using K-type thermocouples attached to each end of the sample.Results and DiscussionFactors of interest in terms of weld bead geometry were the build-up height per pass (H) and average wall thickness (W) as well as the apparent surface roughness of the build-up which was characterised in terms of machining layer thickness (MLT). This was defined as the1017maximum thickness of material required to be removed (by a machining or similar process) to bring the deposit back to a solid continuous vertical surface.

This was defined as the1017maximum thickness of material required to be removed (by a machining or similar process) to bring the deposit back to a solid continuous vertical surface. This was used in preference to other measures such as Ra or standard deviations since it relates directly to the amount of tool engagement required in finishing processes and also represents potential material wastage. Height per pass was measured at the completion of each pass using a Micro-Epsilon scanCONTROL 2710-50 laser profile scanner held adjacent to the welding torch. Additionally, the sides of each sample were profiled at the completion of each experiment, with measurements used to determine wall and machining layer thicknesses. These geometry factors, illustrated in Figure 1, were used as output variables in the modelling process. Deposition rates (DR) were also determined for each experiment with knowledge of both the wire feed speed and actual mass deposited.

Deposition rates (DR) were also determined for each experiment with knowledge of both the wire feed speed and actual mass deposited. These results are summarised in Table 4.(a)(b)Figure 1. Sample geometry with (a) coordinate system and (b) geometry factors defined.Table 4.

As such, interpass temperature was not considered as a factor in all subsequent analyses and modelling. The F-statistics results from 4-way analysis of variance (neglecting interpass temperature) for the weld deposit geometric outputs of interest are summarised in Table 5.Table 5. F-statistic values from 4-way ANOVA.Input ParametersHOutput Variables WMLTL I WFSR TS4.32 18.53 59.01 16.450.43 10.88 0.19 11.720.57 0.47 0.94 1.09Considering firstly the average wall thickness, W, it can be seen from the comparatively large F-statistics values for arc current and travel speed that these two input parameters appear to have the greatest influence over the thickness of the weld deposit. This may be readily explained as both arc current and travel speed, when combined with arc voltage, relate directly to arc energy and so heat input.

This may be readily explained as both arc current and travel speed, when combined with arc voltage, relate directly to arc energy and so heat input. As heat input is a measure of the amount of thermal energy delivered to the weldment, it can be clearly understood that for a given heat input, a molten pool of a given size will be created (assuming other factors such as the thermal properties of the weldment are consistent, as in the present study). The width of the molten pool then acts as to limit the width of the deposited bead, and so the thickness of the wall type build-up that forms the basis of this additive manufacturing technique. This finding is consistent with the models for weld bead width during conventional GTA butt weld developed by of Esme et.al. [7] and Tarng and Yang [9].In contrast, the ANOVA results for build-up height per pass show a strong dependence on wire feed speed, with arc current and travel speed having a lesser influence.

[7] and Tarng and Yang [9].In contrast, the ANOVA results for build-up height per pass show a strong dependence on wire feed speed, with arc current and travel speed having a lesser influence. This observation is self-evident as wire feed speed is directly related to the material deposition rate during the welding process. Given that the width of the deposit is effectively controlled by heat input, conservation of volume infers that the addition of filler material will yield a proportional build- up in terms of height per pass.

It is this independence of wire feed from heat input that gives GTA welding a considerable advantage over GMA and other arc-wire based welding processes, offering precise control over bead geometry and surface quality, and so making it well suited to additive manufacture methods [11].While not quantified, results of this study indicate that there exists a physical limit to the deposition rate possible for a given heat input while maintaining acceptable weld quality. This was particularly evident in experiment four where the combined low heat input and high wire feed speed produced the ‘stubbing’ of the wire on the solid-liquid interface within the molten pool, as seen in Figure 2 by the deposition of ‘folded’ sections of unmelted filler material. By comparison, the sidewall profile of experiment eight, which was deposited using a more realistic parameter set, is significantly more desirable.

By comparison, the sidewall profile of experiment eight, which was deposited using a more realistic parameter set, is significantly more desirable. Despite such qualitative observations, the results of ANOVA presented in Table 5 show no clear correlation between the chosen measure of1019surface roughness and the input parameters considered. It is considered that the MLT may be a function of the thickness-to-height ratio as well as heat input and deposition rate however further investigation is required to test these hypotheses.Figure 2. Samples from experiments four and eight showing variation in sidewall surface profiles.Using the experimental results obtained, mathematical models relating input parameters and output variables were developed using empirical curvilinear equations [12]. Here, geometrical output variables, represented by the response parameter Y, are related to the four key input parameters under the assumption of linear relationships for close ranges.

Here, geometrical output variables, represented by the response parameter Y, are related to the four key input parameters under the assumption of linear relationships for close ranges. This yielded equations of the form:where:Y is any output variable such as H, W or MLT X1-X4 are the input parameters of L, I, WFSR and TS respectively a, b, c, d and k are modelling constantsValues of modelling constants for each geometrical output were determined using multiple regression techniques, with these results presented in the following equations.

Within these models, the relative influence of each input parameter may be determined by consideration of the power to which it is raised.Build-up height per pass:Average wall thickness:Thickness-to-height ratio:Machining layer thickness:As a means of comparison, values for each geometric output were calculated using the mathematical models derived and then plotted against the corresponding experimental data as shown in the scatter plots of Figure 3.1020(1)(2)(3)(4)(5)(a)(b)(c)(d)Figure 3.

Comparison of measured and calculated values for (a) build-up height per pass, (b) average wall thickness, (c) thickness-to-height ratio and (d) machining layer thickness.The relative error, e, associated with each calculation from mathematical models wasdetermined using:||(6)where:m is the measured value for the geometrical factor of interest, and c is the calculated value for the same geometrical factor.The errors accompanying each of the mathematical models are summarised in the histogram presented in Figure 5. This illustrates that models used to relate build-up height and wall thickness to the selected input parameters are in close agreement with experimental data with an average accuracy of 91% and 95% respectively. The model generated to predict thickness-to-height ratios of build-up deposits still shows good agreement with experimental data with an average accuracy of 88%.

The model generated to predict thickness-to-height ratios of build-up deposits still shows good agreement with experimental data with an average accuracy of 88%. In contrast the model for machining layer thickness has poor correlation to the experimental data with an average accuracy of 68%. This low level of accuracy is to be expected given the apparent lack of influence of the selected input parameters. As such, further work in this area is required in this area, be it the investigation of relations to dependent variables such as heat input, deposition rate and thickness-to-height ration, or the definition of a new measure for sidewall surface roughness.1021108t n u o C6Height, H4Thickness, WW:H ratio2MLT00 - 5%5 - 10% 10-20% Magnitude of Error>20%Figure 4.

Accuracy analysis of models showing errors between measured and calculated values.It is hoped that empirical relations, such as those developed in the present study, may be used to optimise the GTA welding process for additive manufacture and so allow the process to be competitive on a cost basis with more conventional manufacturing methods, particularly for high cost components such as those fabricated from titanium alloys for use in the aerospace sector. Escobar-Palafox et.al. [13] have also shown that process models can be used to generate control regimes allowing for more flexible automation of the GTA-wire based additive manufacturing process.ConclusionsGas tungsten arc welding has been used in conjunction with ‘cold’ wire feed to successfully generate additively manufactured profiles in Ti-6Al-4V. Using Taguchi DOE techniques, experiments were conducted to investigate the influence of common welding process parameters on the geometry of the resulting multi-layer deposit.

Using Taguchi DOE techniques, experiments were conducted to investigate the influence of common welding process parameters on the geometry of the resulting multi-layer deposit. Results show that the average thickness of the wall type deposit is primarily a function of arc current and travel speed while build-up height is chiefly related to wire feed speed, and hence deposition rate. Empirical curvilinear models relating the process parameters considered to the geometrical factors of interest were generated using multiple regression techniques, with models for wall thickness and build-up height showing good agreement with experimental data. Conversely, machining layer thickness, used to characterise sidewall surface roughness, showed very little dependence on any of the welding process parameters considered, with this being reflected in the relatively poor accuracy of the mathematical model.

The mathematical models developed show considerable promise for both process optimisation and control of process automation.AcknowledgementsThe authors would like to acknowledge the support of the Defence Materials Technology Centre (DMTC). The DMTC is supported by the Defence Materiel Organisation. In addition they would like to acknowledge the valuable contributions of Dr Zengxi Pan and Mr Nathan Larkin as well as the supervision of Professor John Norrish and Professor Rian Dippenaar.1022References1.2.3.4.5.6.7.8.9.10.Seong, S., O. Younossi, and B.W. Goldsmith, Titanium: Industrial Base, Price Trends, and Technology Initiatives. 2009, Santa Monica, California: RAND Corporation. 155. Kraft, E., Summary of Emerging Titanium Cost Reduction Technologies. 2004, EHK Technologies-for ORNL: Vancouver, WA. p. 1-59. Benyounis, K.Y. and A.G. Olabi, Optimization of different welding processes using statistical and numerical approaches – A reference guide.

E CHARACTERISTICS AND APPLICATIONS OF CERAMIC LASER FUSIONAND CERAMIC LASER SINTERING H. H. Tang1, H. C. Yen1,*, M. L. Chiu2, J. H. Jou31 Department of Mechanical Engineering, National Taipei University of Technology 2 Graduate Institute of Mechanical and Electrical Technology, National Taipei University of Technology 3 Graduate Institute of Manufacturing Technology, National Taipei University of Technology *Corresponding author Reviewed, accepted September 10, 2008AbstractThe aim of present study is to investigate the possible application of the ceramic partswhich are fabricated with the process of Ceramic Laser Fusion or Ceramic Laser Sintering.The experimental results reveal: (1) CLF can lead to a reduction in the porosity of theceramic part but also can induce micro-cracks. Therefore, this process cannot produce a partwith the required strength by a post-process of infiltration; (2) CLS is capable of fabricating aceramic part with high porosity.

Therefore, this process cannot produce a partwith the required strength by a post-process of infiltration; (2) CLS is capable of fabricating aceramic part with high porosity. By adjusting the slurry formulation and varying the scanningenergy, the open porosity can be over 90vol% of the total porosity. After a post-process ofinfiltration, the density can be increased to 95%; therefore, CLS can apply to produce a partwith high strength. Because the high open porosity leads to a good permeability, the processof CLS is suitable for the fabrication of ceramic shell mold.Keywords: Ceramic Laser Fusion; Ceramic Laser Sintering; porosity; ceramic shell mold1. IntroductionRapid Prototyping is a non-traditional technology based on layer manufacturing.

IntroductionRapid Prototyping is a non-traditional technology based on layer manufacturing. To date,some processes have been commercialized, such as Stereolithography (SL)[1,2], SelectiveLaser Sintering (SLS)[3,4], Laminated Object Manufacturing (LOM)[5,6], and 3D Printing(3DP)[7,8]; all these techniques have the potential of fabricating ceramic parts. Owing to thedemand of producing ceramic functional parts, some researches are switched to investigating the feasibility of fabricating ceramic components based on those processes.Tang[9,10] invented two rapid prototyping processes, Ceramic Laser Fusion (CLF) and Ceramic Laser Sintering (CLS), in 2001 and 2007 respectively. Both processes are somewhat similar to SLS in principle of using laser scanning to fabricate parts; nevertheless, instead of using loose powder in SLS, slurry which consists of ceramic powder and inorganic binders is employed in CLF/CLS.

The ceramic slurry allows the use of smaller ceramic particles to castultra-thin layer which leads to a higher green density. Both processes have identicalmanufacturing steps but different working temperature. During the laser scanning, theworking temperature of CLF is higher than the melting point of the ‘high melting point’ceramic powder; because the high melting point ceramic powder is melted, the consolidationmechanism is defined as ‘fusion’ in this paper. On the other hand, the working temperature ofCLS is lower than the melting point of the high melting point ceramic powder, but higherthan the melting point of the low melting point ceramic powder; consequently, the highmelting point ceramic powder remains solid while the molten low melting point ceramicpowder spreads between the solid particles almost instantaneously as it is driven by intensecapillary forces; the binding mechanism is defined as ‘sintering’ in this paper.

The ‘highmelting point’ ceramic powder is so-called the structural material, while the ‘low meltingpoint’ material is then what is called the binder.The part density is correlated with the working temperature. Actually, CLF and CLShave a different working temperature to each other. By comparison with CLS, the partsfabricated by CLF can achieve a higher density and strength without any post-process ofdensification. The parts fabricated by these two processes demonstrate different structurecharacteristics to one another. The aim of current study is to investigate characteristics andapplications of the parts made by these two processes. The slurry based rapid prototypingtechnology with controlled laser energy was utilized to fabricate ceramic parts for measuringthe part porosity and evaluating the part’s mechanical characteristics. Eventually, the feasibility of both processes for producing the parts with high or low mechanical strength was evaluated.2.

Eventually, the feasibility of both processes for producing the parts with high or low mechanical strength was evaluated.2. Experimental procedureTechnology for fabricating ceramic parts by CLF is well established. The slurryformulation, consolidation mechanism, and laser scanning parameters for CLS wereinvestigated by Yen[11]; a knowledge base was built to establish the relationship among thescanning line width, scanning line depth, consolidation mechanism, and working parameters;it is in favor of selecting parameters for experiments.To fabricate a 3D part, fabricating a specimen (Fig. 1) to obtain scanning line width and depth in advance is a must. A layer with a thickness of 10 (cid:2)m for observing the scanning depth was fabricated on a pre-fabricated solid base. Afterward, some spacers were built on the top of observing layer, and another layer with several separated scanning lines for observing the scanning line width was fabricated on the top of spacers.

Afterward, some spacers were built on the top of observing layer, and another layer with several separated scanning lines for observing the scanning line width was fabricated on the top of spacers. After the process of green part removal, the depth of overhang, which is the scanning line depth, and the scanning line width were measured by an optical microscope equipped with an X-Y platform driven bymicrometer head. Based on the collected data, the scanning space and the layer thickness for fabricating a full melting part can be achieved. A CO l2 aser has been used in this experiment. Scanning energy density is varied with laser beam size, laser power and scanning speed;selecting a suitable parameter combination to achieve a high energy density laser scanning isthe principle of building a fusion specimen. After the green part removal, the consolidationmechanism can be verified by observing the topography with a Scanning ElectronMicroscope (SEM).Fig.

After the green part removal, the consolidationmechanism can be verified by observing the topography with a Scanning ElectronMicroscope (SEM).Fig. 1 Schematic of fabricating a specimen for measuring scanning line width and depth2.1 Fused specimens fabricated by Ceramic Laser FusionIn the present study, the material was prepared with the proportion of 52.63 wt% silica, 3.68 wt% clay, 1.58 wt% silica sol, and 42.11 wt% water. Particle size of silica is 11~15 (cid:1268)m. The well-dispersed slurry was obtained by mixing the slurry with an electric blender for 12hours. A 10 x 10 mm fused specimen was made by scanning a laser beam over the surface ofthe green layer. The layer thickness was 40 (cid:2)m.

A 10 x 10 mm fused specimen was made by scanning a laser beam over the surface ofthe green layer. The layer thickness was 40 (cid:2)m. The laser was operated at laser power of19W with a scan speed of 60 mm/s and scanning space of 0.1 mm; this scanning parametercombination led to a linear overlap of 80%, and a planar overlap of 80%.2.2 Sintered specimens fabricated by Ceramic Laser SinteringTwo different compositions of the slurry were used as shown in Table 1. Slurry I contained more binders than slurry II did. Layer with a thickness of 20 (cid:2)m was casted on a working platform which was preheated to 90 (cid:1082). Following the same process as mentioned in section 2.1, the specimens were fabricated for measuring the scanning line width and depth.

Following the same process as mentioned in section 2.1, the specimens were fabricated for measuring the scanning line width and depth. By collaborating with a variety of laser power, a 10 x 10 mm sintered specimens was fabricated with a parameter combination as following: scanning speed of 500 mm/s, laser power of 14W, scanning spacing of 0.15 mm, layer thickness of 20(cid:1268)m, linear overlap of 75%, planar overlap of 55%; and de-focus of 4.75 mm.Table 1 The composition of the slurry for fabricating ceramic sintering specimensSiO2 (g) Clay (g) Silica sol (g)Slurry I 100 7 9Slurry II 100 3 72.3 Porosity measurementThe liquefied wax was infiltrated into the open pores between the solid by capillaryforces. Equations 1~6 can be used to individually figure out bulk density, open porosity and closed porosity of the part. (cid:568)A(cid:569)represents the weight of a laser sintered part which was measured after cooling at room temperature for an hour.

(cid:568)A(cid:569)represents the weight of a laser sintered part which was measured after cooling at room temperature for an hour. The specimen was immersed in the 70 (cid:666) liquefied wax for 30 minutes, and then taken out for cooling at room temperature for 30 minutes.(cid:691)(cid:984)B(cid:569)represents the weight of the infiltrated specimen. (cid:568)C(cid:569)represents the weight of the infiltrated specimen which was weighted during it was immersed in a tank filled up with 25 (cid:666)(cid:33)water.

Results and discussionThermal energy delivered from a laser is absorbed by the green part, and is transferred from the surface to the interior. Fig. 2 schematically shows a temperature distribution corresponding to the energy distribution of the Gaussian beam which has the maximum energy intensity at the beam midpoint. Under the condition of constant scanning speed andbeam size, as the laser power is raised to a certain value, the scanned surface will reach the melting point of the clay (~1200(cid:666)), the melted clay binds the solid silica particles to producea liquid phase sintering zone. If the laser power is continuously increased, the silica particleson the surface will be melted, and the fusion zone will extend from the surface to the interiorof the green part.

If the laser power is continuously increased, the silica particleson the surface will be melted, and the fusion zone will extend from the surface to the interiorof the green part. The area below the fusion zone is the sintering zone as mentioned above.The property transformation zone, which includes the fusion zone and sintering zone, can be taken apart from the green block by a green part removal operation. Dpt and WL represent the scanning line depth and width respectively. If the layer thickness is smaller than the fusionzone depth, the fusion zone of each layer will be overlapped; then the fusion part can bemanufactured.Fig. 2 Temperature distribution corresponding to the energy distribution of the Gaussianbeam3.1 The characteristics of the fusion partThe topography shown in Fig. 3 is the surface of the fusion part fabricated with theparameter combination mentioned in sec. 2.1. The surface is not smooth and has manydepressions.

3 is the surface of the fusion part fabricated with theparameter combination mentioned in sec. 2.1. The surface is not smooth and has manydepressions. A few closed pores are also included in the interior of the fusion part; however,the porosity is lower than that of the sintering part. Moreover, cracks are brought out alongthe direction perpendicular to the scanning line. The full fusion parts made with otherscanning parameters have the same results. Consequently, high density, rough surface, andclosed pores are the characteristics of the fusion ceramic part.Fig. 3 The topography of the fusion partThe fusion part provides a higher density than the sintering part does and can result ahigher mechanical strength; however, the high density is resulted from the full meltingceramic powder, but the fusion part contains a large number of micro-cracks which is evenlydistributed throughout the part and deteriorates the strength of the fusion part.

As the ceramicis a brittle material, its tensile strength is only one tenth of its compressive strength and it canhardly deform plastically; therefore, thermal stress will easily induce micro-cracks. In currentinvestigation, in order to achieve full melting, a high laser power with a low scanning speedwas employed.

In currentinvestigation, in order to achieve full melting, a high laser power with a low scanning speedwas employed. Due to the continuous scanning, the surface temperature increases rapidly;therefore, the thermal expansion on the surface is larger than that at the interior of the part.The surface and the interior of the part bear compressive stress and tensile stress respectively;after scanning, the scanned surface contacts the cold air in the atmosphere; therefore, thecooling speed on the surface is faster than that in the interior while the stress direction isopposite to the trend of temperature rising; such phenomenon easily leads to crack formationduring the manufacturing process.3.2 The characteristics of the sintering partSome specimens for measuring the scanning line width and depth were fabricated withthe parameter combination and slurry I shown in section 2.2. As a result, the averagescanning line width was 0.49 mm and depth was 0.045 mm.

As a result, the averagescanning line width was 0.49 mm and depth was 0.045 mm. Based on these data, a part can be fabricated with sintering consolidation mechanism. Fig. 4 shows some 10×10×1 (mm(cid:529) ceramic sintering parts. Fig. 5(a) and 5(c) demonstrate their surface and cross-section. Fig.5(a) shows that sintering consolidation appears on the surface. Although a partial meltingoccurs locally, the individual molten area is less than 20×20 (cid:2)m and does not connect witheach other. Fig. 5(c) reveals the sintering part is porous and no extensive melting area, whichprohibits the connection of the pores, exists between the layers.The surface of the part shows more melting state when the laser power was increased to16W. Fig. 5(b) and 5(d) demonstrate their surface and cross-section. The melted silica fills up the pores on the surface. An extreme thin fusion layer exists between the two layers as shown in Fig.

Fig. 5(b) and 5(d) demonstrate their surface and cross-section. The melted silica fills up the pores on the surface. An extreme thin fusion layer exists between the two layers as shown in Fig. 5(d); such fusion layer obstructs the interconnection of the pores on the boundary of the layers, and then reduces the amount of open pores dramatically. By collaborating with a variety of laser power, de-focus distance of 4.75 mm, and scanning speed of 500 mm/s, the porous sintering parts also can be fabricated with the slurry II which contains less binder as mentioned in section 2.2.Fig. 4 Parts (10 x 10 mm) fabricated by the process of Ceramic Laser sintering(a)(a)(b)(d)Fig.

4 Parts (10 x 10 mm) fabricated by the process of Ceramic Laser sintering(a)(a)(b)(d)Fig. 5 (a) The topography of the part built with a laser power of 14 Watts(b) The topography of the part built with a laser power of 16 Watts(c) The cross section of the part built with a laser power of 14 Watts(d) The cross section of the part built with a laser power of 16 Watts3.3 The porosity of the ceramic sintering partsTable 2 shows the porosity of the parts fabricated with different laser power and twokinds of slurry mentioned in section 2.2. The total porosity of the sintering part made ofslurry I is 46%, but the closed porosity is higher than that in the part made of slurry II. If alow melting point material infiltrates into such parts, the highest density is only 84%. Thiscan be explained that the molten clay blocks up a part of pores to form closed pores duringthe process of binder melting so that the number of the closed pores is increased.

Thiscan be explained that the molten clay blocks up a part of pores to form closed pores duringthe process of binder melting so that the number of the closed pores is increased. Owing tothe higher proportion of the binder, the closed porosity of the part fabricated with the slurry Iis higher than that of the part fabricated with the slurry II.Table 2 The porosity of the parts fabricated with different laser powers and two kinds of slurry listed in Table 1Laser Power (W)Open porosity (cid:528)%(cid:529)Closed porosity (cid:528)%(cid:529)Slurry I14301610387.8Slurry II810 (with post-sintering)3740.67.45.28 (with post-sintering)42.54.1For the slurry II, a minor adjustment of the laser power can not lead to an obviousvariation of the total porosity.

The total porosity of the slurry II is almost the same as theresult of the slurry I; however, because the proportion of the binder is only 10vol% of the total ceramic materials, the closed porosity of the slurry II is reduced to 7~8(cid:664) while the open porosity is increased to38%. When a low melting point material infiltrates into thesintering part, the density is increased to 92%. Furthermore, when the sintering part is put into a furnace with a constant temperature 1150 (cid:666) for 2 hours, the total porosity is almost not changed, but the closed porosity is reduced around 3(cid:664) while the open porosity is increased to 40 ~ 42% and the density can be increased to 95% if it is infiltrated afterwards.Fig. 6(a) is the cross section of the part which is fabricated with a laser power of 8W with athermal post-sintering; Fig. 6(b) is the cross section of the part infiltrated with wax.Fig.

6(a) is the cross section of the part which is fabricated with a laser power of 8W with athermal post-sintering; Fig. 6(b) is the cross section of the part infiltrated with wax.Fig. 6 (a) The cross section of the part which was fabricated with a laser power 8W and aprocess of post-sintering; (b) the cross section of the part infiltrated with wax4. ConclusionAccording to the characteristics of the fusion part and the sintering part mentionedabove, we can evaluate the feasibility of applying these two methods for producing parts withhigh mechanical strength and parts with low mechanical strength.In comparison with CLS, CLF can lead to a reduction in porosity and a higher density. Because silica is a brittle material, a high energy laser scanning easily leads to thermal crack on the part; such cracking might bring a minor defective to a destructive crack development.

Because silica is a brittle material, a high energy laser scanning easily leads to thermal crack on the part; such cracking might bring a minor defective to a destructive crack development. Because a great portion of the pores in the fused structure are closed, the molten material can not be infiltrated into the pores of the fused structure; therefore, the mechanical strength of the part can not be increased by infiltration. Obviously, CLF is also not suitable for the part which needs high permeability, such as ceramic shell mold for precision casting.CLS can be used to fabricate the part with high porosity. By changing the slurryformulation and varying the scanning energy, the open porosity can be adjusted. The resultsof the experiments demonstrate that the total porosity of a sintering part can be greater than45% and the open porosity can be greater than 90% of the total porosity.

The resultsof the experiments demonstrate that the total porosity of a sintering part can be greater than45% and the open porosity can be greater than 90% of the total porosity. After infiltrating, thedensity of the part can increase to 95%; such process can be employed for the part needs highmechanical strength. Besides, the high open porosity leads to the process of CLS is alsosuitable for manufacturing ceramic shell mold which needs high permeability.The future work will concentrate on CLS to fabricate ceramic part with high openporosity for the products need high permeability, such as ceramic filter and ceramic shellmold. For the part needing high mechanical strength, a post-process of infiltrating a highstrength metal or non-metal material could follow the process of CLS.Acknowledgement This work was supported by National Science Council of ROC under projectNSC96-2221-E-027-103-MY2.Reference1.J. W. Halloran, M. Griffith, and T.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCHARACTERIZATION AND ANALYSIS OF GEOMETRIC FEATURES FOR THE WIRE-ARC ADDITIVE PROCESSChristopher Masuo*, Andrzej Nycz*, Mark W. Noakes*, and Lonnie J. Love*Manufacturing Demonstration Facility, Oak Ridge National Laboratory, Knoxville, TN 37932, USAAbstractThe wire-arc additive manufacturing (AM) process expands the possibilities of effectively producing large-scale, complex metal objects through high deposition rates at low costs. However, this process is prone to irregularities in geometric features that occur from improper thermal conditions and build parameters that cause uneven build heights. This paper discusses a method to obtain consistent build characteristics and near net shape geometric features for the wire-arc AM process.

This paper discusses a method to obtain consistent build characteristics and near net shape geometric features for the wire-arc AM process. Process parameters are established for each material printed to ensure characterization of layer build height and even flow in the interior of parts. Various sections of the build including perimeter, infill, and various wall thicknesses require different strategies to correctly build the part. Open-loop build geometry is still not sufficient to build a part to near net shape of the original model. Average layer height is determined and used with adaptive height control to print the correct, modeled height.IntroductionThe wire-arc additive manufacturing (WAAM) process utilizes arc welding (e.g., metal inert gas (MIG), tungsten inert gas (TIG), or plasma arc welding (PAW)) and a gantry or robot system to deposit metal in a layer by layer process [1, 2].

Wire-arc is a direct energy deposition (DED) process with features suitable for medium to large-scale metal printing. Wire-arc is a low cost process both in equipment and material. Welding wire and equipment are inexpensive and easily obtainable. Wire-arc is efficient, environmentally friendly, and a safe process [1-5].Specifically, a large-scale additive manufacturing (AM) system requires a build envelope where the longest axis has a minimum reach of approximately 1-2m [3]. For a large-scale wire-arc additive system, a robot’s high range of motion makes it more suitable for this process and enables the ability to print large parts [6]. Welding robots are also well-known in industry, which requires only slight modifications to an existing industrial automated welding infrastructure.

Welding robots are also well-known in industry, which requires only slight modifications to an existing industrial automated welding infrastructure. By utilizing a welding robot for a wire-arc additive purpose, metal big area additive manufacturing (mBAAM) is made possible.Utilizing mBAAM to create large, complex metal objects requires a suitable process for near net shape quality. The challenges are that process parameters need to be determined as well as understanding the characteristics and geometric features that are capable in an arc-wire additive process. Getting started can be difficult; however, this paper provides a guideline to developing, characterizing, and analyzing features that are used as the building blocks for successfully printing near net shape large metal objects. This paper starts with a system description and details how printing was achieved. The method section discusses the features that were studied.

This paper starts with a system description and details how printing was achieved. The method section discusses the features that were studied. Results areThis manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes. The Department of Energy will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).411then shown, and adaptive control is addressed to make further improvements.

This paper concludes with using these methods to assist in post-process refinement.System and Printing Process BackgroundIn this paper, an mBAAM system (Figure 1) developed by Wolf Robotics and installed at the Manufacturing Demonstration Facility of Oak Ridge National Laboratory was used. This system uses an ABB 2600 six degree of freedom robot equipped with a Power Wave R500 Lincoln Electric MIG welder. It has an effective build envelope of about 2ft length x 3ft width x 7ft height (0.6m x 1.2m x 2.1m) and a deposition rate ranging from 3.5lbs/hr (1.6kg/hr) to 15lbs/hr (6.8kg/hr). The current system has two torches attached to the end-effector of the robot for multi- material capabilities. Only one torch was used for the experiments done in this paper.Figure 1. mBAAM Wolf Arc System with multi-torch setupThe process of printing parts using this system has similarities to the common 3D printing methods. A model is created using a CAD software.

mBAAM Wolf Arc System with multi-torch setupThe process of printing parts using this system has similarities to the common 3D printing methods. A model is created using a CAD software. This model is then saved as an STL to load the object into slicing software. The slicing software used was ORNL Slicer, which was developed by researchers at the Manufacturing Demonstration Facility. The slicer breaks the object into layers of tool paths and generates a g-code file. This g-code is then parsed into a robot path using Wolf Robotics’ application add-on for ABB Robot Studios. This printing process is shown in Figure 2.Figure 2. Printing process used for mBAAM system412MethodsA. Straight WallsBeginning with simple geometry, walls provide the features necessary to understand bead- to-bead forming, surface finish, and geometric dimensions.

Straight WallsBeginning with simple geometry, walls provide the features necessary to understand bead- to-bead forming, surface finish, and geometric dimensions. Before printing a tall wall, welding processes were developed and observed by analyzing the first layer of a single-bead and two-bead thick wall (Figure 3). These welding parameters consist of welding (robot travel) speed, wire-feed speed, and welder specific settings that control the metal transfer. In this case, Lincoln Electric has a waveform library that provides multiple processes. For mild steel printed on the mBAAM system, a short circuit transfer called Surface Tension Transfer (STT) was used for mild steel. STT gives the capability of adjusting heat separately from the wire-feed speed as well as enabling low heat input, reduced of spatter, and better control of the weld bead with increased travel speed [7]. For 410 stainless steel, Power Mode was used for stable arc welding, which is not metal or gas specific.

For 410 stainless steel, Power Mode was used for stable arc welding, which is not metal or gas specific. Power Mode also creates a constant output of power [8].Figure 3. First layer of two-bead thick walls with an L-shape and straight patternAfter analyzing and obtaining welding parameters from the first layer, single-bead and two-bead thick walls were printed. With the single-bead walls, the thickness and height of the bead were measured. These measurements were important for deciding the layer-by-layer dimensions [9]. This data provided a starting bead-to-bead spacing for the two-bead thick walls. When printing a complete two-bead wall, bead-to-bead spacing was adjusted until proper forming between beads was made. Adjustment was achieved by using the ORNL Slicer, which does a center bead-to-bead spacing as shown in Figure 4. Cross-sectional cuts on the walls were done to check for proper forming. Voids appeared when using wide spacing shown in Figure 5.

Cross-sectional cuts on the walls were done to check for proper forming. Voids appeared when using wide spacing shown in Figure 5. By decreasing the bead-to- bead spacing, these voids were diminished. Once a proper bead-to-bead spacing was obtained, the two bead walls were measured in both height and width. Four-bead thick walls were printed and measured afterwards. The printed one, two, and four-bead thick walls are shown in Figure 6.Figure 4. Center bead-to-bead spacing depicted in the ORNL Slicer413Figure 5. Cross-sectional cuts of a two-bead wall with large voidsFigure 6. One, two, and four-bead thick walls printed in 410 Stainless SteelB. Circular TubesAnother simple shape, circular tubes, may represent internal features such as internal cooling passageways for hot stamping dies or engine block cylinders. This shape was also chosen to understand how the flow of material would change, if any, when encountering a circular path.

This shape was also chosen to understand how the flow of material would change, if any, when encountering a circular path. The objective of printing circular tubes was to observe the size change of the inner and outer diameters. For this part of the experiment, one and two-bead thick circular tubes (Figure 7) were printed. The inner diameter of the tube ranged from 8mm to 30mm with 2mm increments. Measurements on the inner and outer diameters were made to determine an inner and outer offset of the circular tubes. Figure 8 shows the representation of the offsets from the modeled tube. These tubes were also scanned using a FARO Arm Laser Scanner to compare the printed parts with the CAD models.Figure 7. One and two bead thick circular tubes with different inner diameters printed in 410 Stainless Steel414Figure 8. Visualization of the inner and outer offset of a circular tubeC.

One and two bead thick circular tubes with different inner diameters printed in 410 Stainless Steel414Figure 8. Visualization of the inner and outer offset of a circular tubeC. Square BlocksLike walls, square blocks have a perimeter and an inset, which are the same as the bead- to-bead spacing shown in Figure 4. In addition to perimeter and insets, infill is introduced when printing a square block. In ORNL Slicer, there are many options for infill patterns (i.e., concentric, line, grid); however, the line pattern was chosen since it decreases the amount of starts and stops when printing as well as provides more heat into the part. When using a line pattern, the linear length infill offset and the sparse infill line distance (Figure 9) were adjusted. The objective was to have flat layers. The right amount of overlapping of beads is necessary to prevent having “hills” and “valleys” in the print.Figure 9.

The objective was to have flat layers. The right amount of overlapping of beads is necessary to prevent having “hills” and “valleys” in the print.Figure 9. ORNL Slicer representation of the types of bead spacing used in the solid blockResultsStarting with walls, Figure 10 shows a cross-sectional cut of a two-bead thick wall with no voids. The material of this wall was 410 stainless steel, and the bead-to-bead spacing used was 3.2mm. The measurement data obtained from these walls is shown in Table 1. Due to running the first layer with a slower weld speed to heat up the build plate, the first layer’s height and width (thickness) was measured in addition to measuring the wall when it was completed. In Table 1, a height per layer of the walls was determined to input back into the slicer, which ensures that a part415will print to the correct height. This value will also be used with adaptive, closed-loop control to minimize the error in height when printing.Figure 10.

This value will also be used with adaptive, closed-loop control to minimize the error in height when printing.Figure 10. 410 stainless steel, two-bead thick wall without any voidsTable 1. Average measurement data from a one, two, and four- bead thick stainless-steel wallFigure 11 shows a graph of the printed inner and outer diameters and offsets of the circular tube compared to the CAD model, and Figure 12 shows the scanned tubes and CAD model comparison. In Figure 11a, the comparison between the actual and designed tubes have a linear relationship for both the inner and outer diameters. There was not a drastic difference in the inner and outer diameters between the one and two-bead thick tubes. In Figure 11b, the inner offset was larger than the outer offset by approximately 1.5mm. This indicates that the inner diameter shrinks more than the outer diameter.

In Figure 11b, the inner offset was larger than the outer offset by approximately 1.5mm. This indicates that the inner diameter shrinks more than the outer diameter. When designing features such as cooling channels, the CAD model’s bore features should be larger to avoid the cooling channels from closing mid print.416Figure 11. Comparison of the as designed tubes to the actual tubes in 410 stainless steel. a.) Inner and outer diameter plot. b.) inner and outer offsets vs. the inner diameter of the tubesFigure 12. Scan of the tubes compared with the CAD model. This was done with Geomagic Control X.The results for the square block was more observational. Initial blocks had distinct overbuilding (Figure 12a) in the first few layers. In Figure 12b, a nice, flat square block was printed with 15 layers out of 410 stainless steel. The perimeter and inset value used was 3.2mm, which is the same as the bead-to-bead spacing obtained from the straight walls.

The perimeter and inset value used was 3.2mm, which is the same as the bead-to-bead spacing obtained from the straight walls. The sparse infill line distance used was 4mm with a linear infill length offset of 2.4mm. Due to having one perimeter and one inset, the layer height used was 2.69mm, which was obtained from Table 1. These were successfully printed without adaptive control; however, if the block was significantly taller, the error in the height would accumulate, leading to sloped layers.Figure 13. a.) Block with too large overbuild in just two layers, b.) No overbuild with 15 layers in.These parts were relatively short (40mm tall) and did not experience an uneven build. When printing a part larger than 100mm, overbuilding accumulates leading to uneven surfaces and the possibility for the part failing completely. For large-scale printing, it is crucial to have417consistent layer heights with flat surfaces to achieve a near net shape geometry.

For large-scale printing, it is crucial to have417consistent layer heights with flat surfaces to achieve a near net shape geometry. An open-loop control system is not able to do this, which makes a closed-loop system necessary [10].Wire-arc additive systems that use a closed-loop control system tend to use vision-based sensors and process this information to adapt for the next layer. In the mBAAM Wolf Arc system, closed-loop control is achieved by using wire tracking. Wire tracking measures the impedance of the welding arc and corrects the wire stick-out (Figure 14) based on a specified current [11]. By having the end of the stick-out to be the tool center point (TCP) of the robot, the height of the build can be determined during metal deposition. By combining both feedback and feedforward control, adaptive height control can be achieved using minimal hardware. This also results in having height correction during deposition, creating flat, smooth layers.

This also results in having height correction during deposition, creating flat, smooth layers. The adaptive height control process is shown in Figure 15. Basically, the welding speed of the robot either slows down or speeds up depending on whether there is an overbuilt or underbuilt surface. Wire-feed speed is held constant, allowing metal to deposit more or less depending on the surface height. Drastic improvements can be seen in tall walls shown in Figure 16.By combining the geometries used in this paper, the process of printing a complex shape was simplified. Figure 17 shows the integration of all these geometries. This part was a sample hot stamping tool with integrated cooling channels printed using 410 stainless steel. Using adaptive control also resulted in a smooth, flat surface. The final result led to a more near net shape geometry.Figure 14. MIG wire stick-out, which is the distance between the end of the contact tip and the base plateFigure 15.

The final result led to a more near net shape geometry.Figure 14. MIG wire stick-out, which is the distance between the end of the contact tip and the base plateFigure 15. Drawing of the adaptive control process418Figure 16. Comparison between a non-adaptive printed wall and an adaptive printed wallFigure 17. Hot stamping tool with sliced image on the left and printed part on the rightConclusionWith simple geometries, characterization of wire-arc features was broken down into a simple process. By understanding these geometries, complex shapes with near net shape geometries can be developed and achieved with ease. The obtained data also assists with post- process machining, which is used to refine the printed part to produce the actual designed part with minimal material waste.AcknowledgementsThis material is based upon work supported by the U.S.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCHARACTERIZATION AND TESTING OF 3D PRINTED SPROCKETSRohit K. Srinivasan Chap Research, Westlake High School, Austin TX 78746 USAAbstractAlthough arguably only 30 years old, 3D printing is already having a tremendous impact upon a broad spectrum of industries, from medical products to consumer goods and nearly every industry in between1. The primary driver for growth has been the ability to rapidly prototype components at very low cost. However, as 3D printing technology has matured, industry participants are experimenting with printing components for production and long-term use as opposed to just prototypes.Chap Research has created a web-based application used to 3D print a “sprocket” – a toothed wheel used in chain-driven systems.

This paper details our study to determine the applicability of these plastic 3D printed sprockets in production and for long-term use within certain robotics applications, an area traditionally dominated by metal (aluminum) sprockets.The study varied a range of design parameters and evaluated whether plastic (PLA) 3D printed sprockets could withstand the torsional stresses and fatigue failure modes present in robotics applications (specifically for FIRST2 competition robots). This paper also describes two custom test rigs that were built, one to characterize torsional stress and the other for fatigue stress analysis.Our study indicated that, for our target application, FIRST robotics competitions, 3D printed sprockets were quite sufficient in production and for long-term use. Layer height and infill settings for the 3D prints had the largest impact on the performance of these sprockets.

Layer height and infill settings for the 3D prints had the largest impact on the performance of these sprockets. While this is not a study to comprehensively compare performance of 3D printed sprockets to their aluminum counterparts, the big surprise was that 3D printed sprockets performed even better than metal sprockets for fatigue stress modes of failure.

Using these results, this study also makes some recommendations on infill and layer height settings for achieving the desired performance for FIRST robotics applications.The study also showed the relationship between the load applied to a 3D printed sprocket and the number of cycles of continuous operation before a fatigue load related failure occurred and makes a conjecture on a range of load for “infinite life” of operation.Future extensions of this study may focus on additional fatigue loading tests, expanding to broader sets of load and printer conditions, further optimization of build parameters, varying printing materials, and different 3D printing technologies such as SLS.1768A. Background1. About SprocketsA sprocket is a toothed wheel that is used with a chain. Sprockets are typically used to drive two parallel axles with one motor and are one of the important components of drive train design3.

Sprockets are typically used to drive two parallel axles with one motor and are one of the important components of drive train design3. Sprockets vary in size and diameter depending on the target application and the type of chain used. Today most sprockets are made of metal and are mass produced in pre- determined shapes and sizes. Customized metal sprockets can also be made using hobbing tools or using a mill.Figure 1: Typical sprocket.2. Metal SprocketsToday’s manufacturing methods reliably produce metal sprockets that have high dimensional accuracy, high strength, and durability and are usable across a wide variety of applications. If one needs a sprocket with any customization or small variants in the design (as is common in design of custom- built robots), the sprockets have to be custom ordered. Such sprockets are typically expensive and also have large lead times for delivery.

Such sprockets are typically expensive and also have large lead times for delivery. This leads to slow prototyping speed and typically limits the creativity of those designing systems while increasing expense.3. 3D Printed Sprockets3D printing has revolutionized many industries, recently expanding into industrial applications. It provides the ability to prototype quickly and cheaply. 3D printed sprockets look to bring similar benefits to the design and manufacturing of sprockets to robotics and industrial applications. 3D printed sprockets are cheap, easy to create and allow for fast prototyping. Despite these advantages, 3D printed sprockets are relatively new and have not yet seen widespread adoption especially in production or load bearing applications.

Despite these advantages, 3D printed sprockets are relatively new and have not yet seen widespread adoption especially in production or load bearing applications. The primary inhibitors are:i) ii)Lack of design expertise to create 3D printed sprockets Concerns about the strength and durability of 3D printed sprockets beyond simple prototypes, especially given they are predominantly made from plastic4. Key Drivers for Widespread Adoption of 3D Printed Sprockets Chap Research has developed a program called SprocketR4 to rapidly and efficiently design sprockets for specific applications. The program generates STL5 files ready for 3D printing based on user input on the size, number of teeth, chain tension, and sprocket hub type.

The program generates STL5 files ready for 3D printing based on user input on the size, number of teeth, chain tension, and sprocket hub type. Due to the general availability of this application and the benefits of 3D printed sprockets, we hope that custom sprockets grow more popular when building robots for FIRST robotics competitions.However, robot designers need to be confident about the durability and performance of these sprockets in their robots. They will adopt 3D printed sprockets only if these sprockets can perform as well as metal sprockets across a range of loads and not fatigue or incur damage over sustained operation.

They will adopt 3D printed sprockets only if these sprockets can perform as well as metal sprockets across a range of loads and not fatigue or incur damage over sustained operation. Future technologies such as carbon fibers can alleviate this concern, but it could be many years before they are full commercialized.This paper describes our studies to characterize the durability of currently widely availableplastic/polymer based 3D printed sprockets for specific robotics applications.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161769B. Related WorkOur motivation to perform this study was driven by minimal prior art in the arena of characterizing torsional and fatigue stresses in 3D printed sprockets. As per Shigley and Mischke6, Wilfred Lewis introduced an equation in 1892 for estimating the bending stress in sprocket teeth.

As per Shigley and Mischke6, Wilfred Lewis introduced an equation in 1892 for estimating the bending stress in sprocket teeth. This equation remains the basis for most metal sprocket design today and is used in the later sections of this study for characterizing the bending stress in 3D printed sprockets. Others have extended the research to deeper work on the analysis of stress and force distributions in metal sprockets. Marshek7 presented a spring model analysis for the load distribution in chains and in sprocket teeth and compares the results with those from a traditional force-balance analysis. Naji8 provided some findings on analysis of load distribution under torsional load for metal sprockets.With the greater popularity of 3D printing and their usage in load bearing applications, researchers have begun to make early forays into characterizing them similar to the work done on their metal-based counterparts in the past century.

However, studies such as Liu9 focus more on the internal structure of the parts and their implications on the strength of these parts. Limited work has been done on specifically characterizing the load bearing capabilities of specialized components such as sprockets.In an informal survey, researchers have pointed to the need for a custom rig in order to provide empirical characterization of 3D sprocket strength. Such rigs have been used in in the past to test moving parts in the field of research medicine. Baums10 and Tavakolan11 have all described rigs to analyze the load bearing strength of specialized medical devices ultimately targeting limb replacements.

Baums10 and Tavakolan11 have all described rigs to analyze the load bearing strength of specialized medical devices ultimately targeting limb replacements. To our knowledge, none of the papers describe custom rigs in the field of 3D printing which is an area of significant innovation in this study.In summary, while there are many research papers broadly in the stress characterization of metal sprockets and gears, and separately on the strength of 3D printed materials, there has been minimal work in the area at intersection of these two areas. There have also been minimal studies in the description of custom rigs that can be employed for the characterization of 3D printed materials in load bearing applications. This study is perhaps among the first in these areas.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161770C.

This study is perhaps among the first in these areas.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161770C. Operating ParametersThere are a variety of key sprocket design and operating parameters in the real life applications that influenced our study. These parameters include the loads managed, sizes of sprockets needed, and the printer’s manufacturing settings.1. Loads ManagedThe weight of robots in FIRST robotics competitions ranges from 30 lbs. to 120 lbs. In steady state operation, with multiple motors and multiple sprockets, the driving load would be typically divided among many sprockets. However, the torque experienced by a sprocket can be instantaneously quite high when suddenly changing direction, say, going from forward to reverse. Or alternatively, a sprocket might be the key power transfer mechanism to lift a robot off the ground.

Or alternatively, a sprocket might be the key power transfer mechanism to lift a robot off the ground. As a result, these weights are a good proxy for the nature of loads that sprockets could experience in competition. There are three types of load situations depending on the nature of competition or design approach. Standard FIRST Technology Competition (FTC) robot: 30 lbs. High-end FTC robot: 50 lbs. FIRST Robotics Competition (FRC) robot: up to 120 lbs.a) b) c)This study will evaluate the use of sprockets across this spectrum of loads.2. Sprocket Design Size ParametersRobots for FTC fit within an 18” cube, while FRC robots typically fit within a 4’ cube. For these sizes and loads, sprocket dimensions range between 50 and 120 mm in diameter and 16 and 80 teeth. This study will evaluate sprocket strength across these design size values.3.

For these sizes and loads, sprocket dimensions range between 50 and 120 mm in diameter and 16 and 80 teeth. This study will evaluate sprocket strength across these design size values.3. 3D Printing Alternatives Today 3D printed parts can be made using FDM12 (Fused Deposition Molding) printers or SLS13 (Selective Laser Sintering) printers. FDM works on an "additive" principle by laying down heated plastic filament material in layers. Selective Laser Sintering (SLS) is an also an additive manufacturing (AM) technique that uses a laser as the power source to sinter powdered material. FDM printers are significantly cheaper than SLS printers, although the latter can produce more complex shapes and structures. Given the lower cost of FDM printers they dominate the market and are the focus of this study. Within FDM printers, machines could utilize ABS (Acrylonitrile Butadiene Styrene) or PLA14 (Polylactic Acid) as the polymer material.

Within FDM printers, machines could utilize ABS (Acrylonitrile Butadiene Styrene) or PLA14 (Polylactic Acid) as the polymer material. This study only focuses on plastic / polymer parts printed using PLA. A MakerBot Replicator 2X15 is perhaps one of the most popular 3D printers and uses PLA materials and is the printer used in this study. There are obviously parts printed from other materials or technologies that we could study here including ABS or SLS or even futuristic ones such as carbon fibers. Those were not tested and are potential candidates for future studies.4. Key Sprocket 3D Build ParametersThe two most commonly varied 3D build parameters are: a. Layer Height: Layer height is one of the key parameters that affects print quality and speed as it sets the thickness of each layer that is being added to generate a print.

Layer Height: Layer height is one of the key parameters that affects print quality and speed as it sets the thickness of each layer that is being added to generate a print. It is characterized either as the number of layers or the thickness of each layer printed during the additiveCharacterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161771manufacturing process. Greater number of layers is the result of a lower layer height. That in turn will result in greater feature resolution and better adhesion to the build plate. This generally delivers better quality and somewhat contributes to higher strength for the part. However, decreasing the layer thickness also means more layers will have to be printed and the time required for 3D printing the part will be significantly increased.

However, decreasing the layer thickness also means more layers will have to be printed and the time required for 3D printing the part will be significantly increased. One can generally expect lower quality and strength (with all other parameters on the print kept the same) and a faster print with a layer height of over 0.3mm, normal strength and medium time to print with a layer height of 0.2mm and higher strength and slower time to print with layer height of 0.1mm9. In this study, we have ranged the layer height between 0.05mm and 0.29mm. Note that during the study we adjusted the number of layers so that each sprocket is the same width, independent upon layer height.b. Infill: Infill is a value represented in percentage from 0% to 100% that shows how filled a sprocket is with material when printed16.

Infill: Infill is a value represented in percentage from 0% to 100% that shows how filled a sprocket is with material when printed16. At 100% infill, there are no gaps within the printed structure providing maximum strength from the material, but also the greatest material cost, slow speed of printing and greatest weight. About 10% percent infill is common for normal uses. In addition to the infill percentage, the infill “pattern” also has an impact upon strength and potentially affects the symmetry of strength of the printed structure. This study varied infill between 10%, 50% and 100% with a common hexagon infill pattern.5. Failure Modes StudiedWhile there are numerous failure modes, this study will look at two types of failure modes, commonfor sprockets:a. Torsional Failure: Torsion is the twisting of an object due to an applied torque. It is expressed in newton-meters (N·m) or foot-pounds (ft.·lbs.).

Torsional Failure: Torsion is the twisting of an object due to an applied torque. It is expressed in newton-meters (N·m) or foot-pounds (ft.·lbs.). It acts perpendicular to the axis of rotation and along the motion of rotation of the sprocket. The failure that is caused by excessive rotational stresses from torque on a sprocket is called torsional failure and we will characterize it in ft. lbs.b. Fatigue Failure: Fatigue failure is the tendency of a material to fracture by means of repeated alternating or cyclic stresses of an intensity considerably below the torsional failure point. We will look to characterize the fatigue load bearing capacity of a sprocket by measuring the time of continuous operation before failure.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161772D.

Torsional Test AnalysisFor this test we created a set of sprockets and tests with varied parameters in an attempt to determine the effect of those parameters on the torsional strength of the sprocket. Every sprocket created had the same overall thickness – the number of layers was adjusted to accommodate thinner layers. During the test we used #25 chain, so all sprockets were sized to work with that chain.1.

Test ApparatusWe built a test apparatus with the following specifications and layout: a) Ability to load up to 120 lbs. b) Load applied by an arm 1 ft. away from the sprocket axle – all our results will be in ft. lbs. Theload is measured by the scale attached to the load arm.c) Provide tension to ensure the chain doesn’t slip while weight is added.ChainBeam chain is connected toArmChain is fixed here3D printed sprocket1 footScaleChain tension mechanismTurn-buckleFigure 2: Testing rig layout.Figure 3: 3D printed sprocket attached to test apparatus.Figure 4: Test apparatus.Detailed SolidWorks model, materials list and rig construction instructions are provided inAppendix B.3. Test MethodologyAs Figure 2 above illustrates, the arm is tethered to the ground through a turn-buckle. Rotating the turn-buckle slowly draws the arm down and increases the effective torsional load on the sprocket.

Rotating the turn-buckle slowly draws the arm down and increases the effective torsional load on the sprocket. This torsional load is then measured using a digital spring scale. During the study, we rotated the turn- buckle really slowly and noted the increasing torsional load measurements on the scale, until the sprocket failed. The scale attached to the turn buckle times arm length (1 ft.) would indicate how much torque was being placed on the sprocket at the time of failure.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 201617744. Test ResultsThe table below illustrates the torsional load failure torque in ft. lbs. for the sprockets tested and averaged across two independent tests. The failure torque observed in the two tests were between 2% and 8% off each other. The average of their breaking point is below. Separately, we also tested the torsional load failure torque for a comparable sized metal sprocket.

The average of their breaking point is below. Separately, we also tested the torsional load failure torque for a comparable sized metal sprocket. We could not create a breakage even past 140 ft. lbs. of torsional load which was the maximum torque load we could create with our test apparatus.DesignatorDescriptionMean Breaking point (ft.

Conclusionsa) Impact of Infill:Infill appeared to have the most significant impact on sprocket torsional stress failure. The chart below plots the torsional failure load for medium-sized sprockets with 12 layers and infill ranging from 10% to 100%.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 2016177580Mean Torsional failure load (ft. lbs.)70605040S5S23020S40%50%100%150%Infill (%)Figure 5: Torsional failure load vs. infill.b) Impact of layer height:The number of layers positively affects the amount of torque a sprocket could hold although it appears to level off around 20 layers as shown below. This chart plots the torsional failure point for Medium sized sprockets at 100% infill while varying the number of layers (which is the inverse of layer height).Mean Torsional failure load (ft.

lbs. to bend and therefore bend before the sprockets break. As a result, in all instances where a Tetrix axle is used in FTC robots, the sprockets will withstand more torsional load than the axle. Even if Tetrix axles were not used, given that FTC robots weigh 30 lbs. to 60 lbs. and torsional stress points will be less than 6” from the center of the sprocket, torsional loads will generally be well less than 30 ft. lbs. That would mean most 3D printed sprockets will have adequate torsional load performance.ii) This study indicates that for FTC applications a 10% infill for sustained sprocket performance will be sufficient.b) FRC robot implicationsi) The torque that FRC robots are exposed to will likely be less than 60 ft. lbs. (stress points are generally less than 0.5 ft.

lbs. (stress points are generally less than 0.5 ft. away from the center of a robot and according to the rules, the robot MUST weigh less than 120 lbs.).ii) Infill of 100% and a wide spectrum of layer heights have resulted in sprockets that have withstood torsional loads greater than 60 ft. lbs. and will be more than adequate for FRC applications. Even an infill of 50% should be adequate for most use cases given that such sprockets have a torsional load failure of over 60 ft. lbs.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161777E. Fatigue Failure AnalysisFor this test the goal was to create a set of sprockets and load tests with varied parameters todetermine the effect of those parameters on the fatigue strength of the sprocket.1.

Sprocket Design Parameter ValuesWhile there is a vast spectrum of sprockets that can be studied for the effect of fatigue loads, rather than construct a broad experimental design, we decided to start this study with ONE specific sprocket – Large sprocket (diameter – 110mm), 10% infill and 14 layers (0.20mm layer height). The reasons for this choice are as follows:Such a sprocket would be at the lower end of torsional load bearing performance as discussed in the earlier section. This sprocket’s performance under fatigue load would then influence the rest of the study on how to potentially expand to a broader set of sprockets.Such a sprocket will also be used very commonly for both FTC and FRC purposes.2. Load Parameter ValuesA sprocket in a typical FRC robot will be driven by a DC motor that in steady state draws 5A to a peak of 10A of current from a 12V battery source or in other words encounters a 60W to 120W load.

Our plan to analyze fatigue load performance was to assess the time duration of reliable performance of a 3D printed sprocket versus a metal sprocket under such load conditions.3. Test ApparatusThe test apparatus shown in Figure 8 illustrates the test apparatus built for this purpose. The power supply provides a 12V source that drives the Powered Motor which in turn drives the aluminum sprocket. The metal sprocket then drives a 3D printed sprocket which is connected to a back driven motor. The back driven motor serves as a generator (G) and so the 3D printed sprocket effectively generates current. A bank of resistors connected in parallel serve as a load and draw current from this generator and effectively emulate a robot drive load. We set up a bank of 1Ω, 0.5Ω and 0.33Ω resistors from which we could create resistances ranging from 0.167Ω to 1Ω as shown in Figure 9 and in Table 3 below.

We set up a bank of 1Ω, 0.5Ω and 0.33Ω resistors from which we could create resistances ranging from 0.167Ω to 1Ω as shown in Figure 9 and in Table 3 below. The range of resistances was set up to emulate a broader array of load bearing scenarios across which we could then test the relative fatigue stress performance of sprockets.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161778ChainAluminum sprocket3D printed sprocketPowered MotorBack-driven motor / generatorPower SupplyDriven LoadResistor bankFigure 8: Test for fatigue failure analysis.G R1 = 0.33ΩSwitch 1R2 = 0.5ΩSwitch 2R3 = 1ΩSwitch 3Figure 9: Resistor bank configuration that draw current from Generator (G).The resulting resistance (Reffective) can be computed as Reffective.

The resulting1 1 𝑅𝑅2+1 𝑅𝑅1+1 𝑅𝑅3resistance achieved (which can then emulate different load scenarios) by turning on / off these three switches is illustrated in Table 3.=Switch 1Switch 2Switch 3Reffective (Ω)Configuration 1OnOffOff0.330Configuration 2OnOnOff0.199Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161779Configuration 3OffOffOn1.000Configuration 4OnOffOn0.248Configuration 5OffOnOff0.500Configuration 6OffOnOn0.333Configuration 7OnOnOn0.166Table 3: Resistance scenarios from turning on / off the three switches.Lower the resistance should result in higher current draw from the generator and hence the greater power load. Configuration 3 (with the highest resistance) should result in the lowest current ) of the back draw and results in a power generated ( driven motor (generator) to be 40% per VEX17 CIM.

Configuration 3 (with the highest resistance) should result in the lowest current ) of the back draw and results in a power generated ( driven motor (generator) to be 40% per VEX17 CIM. From that the load on the sprocket ( 𝑃𝑃𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷) calculated as:of 32W. We estimate the efficiency (𝜖𝜖𝑃𝑃𝑆𝑆𝑆𝑆𝐷𝐷𝑆𝑆𝑆𝑆𝑆𝑆𝐷𝐷𝑆𝑆)𝑃𝑃𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷 of 80W which is in the mid-range of the 𝜖𝜖 operating range of FRC robots. Other configurations will result in a higher load at the resistors and so 𝑃𝑃𝑆𝑆𝑆𝑆𝐷𝐷𝑆𝑆𝑆𝑆𝑆𝑆𝐷𝐷𝑆𝑆) will emulate an even greater loads than the one typical for mid-range of FRC motor loads.

Other configurations will result in a higher load at the resistors and so 𝑃𝑃𝑆𝑆𝑆𝑆𝐷𝐷𝑆𝑆𝑆𝑆𝑆𝑆𝐷𝐷𝑆𝑆) will emulate an even greater loads than the one typical for mid-range of FRC motor loads. The next section looks at the broader study of fatigue load performance and the impact on the number of cycles of operation.That translates to a load at the sprocket (𝑃𝑃𝑆𝑆𝑆𝑆𝐷𝐷𝑆𝑆𝑆𝑆𝑆𝑆𝐷𝐷𝑆𝑆 =Resistor bank of 1Ω, 0.5Ω and 0.33Ω resistorsFigure 10: 3D Printed sprocket being tested in above testing rig.Detailed SolidWorks model, materials list and rig construction instructions are provided inAppendix C.4. Test MethodologyWe set the resistance to Configuration 3 and connected the power source to the drive motor, allowing it to drive the load motor through the test sprocket. We were visually inspecting the 3D printed sprocket (and the metal sprocket) for wear and tear and measuring the time it took for breakage or operational failure.5.

We were visually inspecting the 3D printed sprocket (and the metal sprocket) for wear and tear and measuring the time it took for breakage or operational failure.5. Test ResultsIn the biggest surprise of this study, after 7 hours and 53 minutes, the metal sprocket failed while the 3D printed sprocket continued to operate. Every one of the metal sprocket’s teeth wore off (Figure 11) while most of the 3D printed sprocket teeth (Figure 12) remained in working condition. This wearing out of the metal sprocket (and the 3D printed sprocket) was quite gradual during the entirety of the fatigue load test.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161780isFigure 11: Metal sprocket post fatigue test.Figure 12: 3D printed sprocket post fatigue test.6. Implications and Recommendations for FTC and FRC RobotsThis is a very encouraging result for the use of 3D printed sprockets for FTC and FRC robots.

Implications and Recommendations for FTC and FRC RobotsThis is a very encouraging result for the use of 3D printed sprockets for FTC and FRC robots. Itsuggests the following:These sprockets will have lesser fatigue load related failures for standard operations even compared to metal sprockets.Given that we performed this test with a sprocket that has a low torsional failure point, we can likely conclude that these sprockets are more likely to fail from torsional stress rather than from fatigue failure.With regard to manufacturing recommendations, a broad range of infill and layer height will likely be acceptable to meet a desired fatigue failure performance.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161781F.

Maximum Bending Stress Characterization The modified Lewis equation for characterizing the bending stress (in a sprocket as provided byASME Y14.5-2009 is as follows:𝜎𝜎)𝑆𝑆Wt= Transmitted load on the sprocket tooth (N) 𝜎𝜎 =𝑊𝑊 𝑃𝑃 𝑌𝑌𝑌𝑌 - the diametral pitch of the sprocket= The circular pitch of the sprocket – 6.52 x 10-2m (in our experiments) = = the width of the sprocket tooth at its base (m) – 2.697 x 10 -3m (in our experiments)𝑝𝑝 𝑃𝑃𝜋𝜋𝑆𝑆𝑌𝑌 Y = 0.47 (as per above ASME standard with teeth count of 60 and ∅ = 25° We will apply this equation to the estimation of the maximum bending stress in the sprocket for the operation of different numbers of cycles of operation and plot the tradeoff between the maximum bending stress that can be applied on the sprocket and the number of cycles of operation that we can achieve from that. In theory, the tradeoff will look as follows.

In theory, the tradeoff will look as follows. The dotted line indicates the theoretical maximum bending stress below which the sprocket will have infinite life or cycles of operation.30) a p M25(s s e r t S g n d n e B x a Mi2015105Max bending stress for infinite life002468101214Number of Cycles (10)^𝑛𝑛Figure 13: Max bending stress vs. Number of cycles of operation of sprocket.1. Estimation of Wt The challenge is the estimation of Wt in the Lewis equation which we accomplished using thefollowing steps.Step 1: Estimate the driven load –.

Estimation of Wt The challenge is the estimation of Wt in the Lewis equation which we accomplished using thefollowing steps.Step 1: Estimate the driven load –. Using a traditional multi meter, we derived the powerconsumed by the resistors.𝑃𝑃𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷Step 2: As described in the previous section, we assumed an overall 40% efficiency () of thegenerator and calculated the load on the sprocket (as:𝜖𝜖𝑃𝑃𝑆𝑆𝑆𝑆𝐷𝐷𝑆𝑆𝑆𝑆𝑆𝑆𝐷𝐷𝑆𝑆)𝑃𝑃𝑆𝑆𝑆𝑆𝐷𝐷𝑆𝑆𝑆𝑆𝑆𝑆𝐷𝐷𝑆𝑆 =𝑃𝑃𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷 𝜖𝜖Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161782Step 3: Compute the angular velocity () of the sprocket in radians per second.

ResultsThe measured RPM at the sprocket and the measured current, voltage and the power across thedifferent load configurations at the resistor banks and averaged across two different sprockets of identical shape, size and gear ratios is provided below.Configurations Configuration 1 Configuration 2 Configuration 3 Configuration 4 Configuration 5 Configuration 6 Configuration 7Reffective (Ω) 0.330 0.199 1.000 0.248 0.500 0.333 0.166𝐬𝐬𝐬𝐬𝐬𝐬𝐬𝐬𝐬𝐬𝐬𝐬𝐬𝐬𝐬𝐬3637 3457 3133 3533 3786 3625 3322𝐑𝐑𝐑𝐑𝐑𝐑Current (A)12.6 10.9 5.6 12.6 9.5 12.6 12.4Voltage (V) 4.4 3.7 5.8 3.8 4.8 4.4 3.555.0 𝑷𝑷𝑫𝑫𝑫𝑫𝑫𝑫𝑫𝑫𝑫𝑫𝑫𝑫 41.0 32.8 51.0 47.0 55.0 44.4(W)Table 4: Power and RPM measurements across different resistance load scenarios.Note that the load encountered at the sprocket is 2.5X the load at the resistors due to the efficiency factor at the generator.

Nevertheless, these observations were counter intuitive that the current drawn and hence the power consumed by the resistors increased only until the resistance decreased to 0.33Ω but then started to decrease even as the resistance decreased further. At this time, we do not have a sound explanation for this phenomenon and could be attributable to the current limitations of the power supply (max current drawn could be greater than the rating of the power supply and so current gets throttled) or the motor / generator current-output characteristics.

However, interpreting this load behavior of these generators, power supply and motors is not in the scope of the study and so we chose to work with Configuration 1 – the configuration that resulted in the greatest load at the resistors and Configuration 3 – the least load on the resistors.We determined the number of cycles to achieve failure as the average of two separate 3D printed sprockets each for Configurations 1 and 3 (four in total) and is reported in Table 5. Here are some of the highlights:Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161783Configuration 1 (55W load) failed on an average after 6 hours and 49 minutes which at 3637 RPM translates to 1,487,533 cycles (3637 RPM x 60 minutes x 6 hours and 49 minutes)Neither of the sprockets on Configuration 3 (32.8W load) failed even after 16 hours of operation and we refer to that as “Infinite Life” – a long duration of operation which is well sufficient for any real-life application.

Note that with 5.5cm radius sprocket, at 3600 RPM and 16 hours of rotation and 1:1 gear ratio between the gear and wheel, it translates to a linear distance traversal of almost 120 kilometers, which for all practical purposes, and particularly for FRC robot related applications, is indeed “infinite life”. 16 hours also translates to about 3.5MM cycles In all these instances, the metal sprocket failed well before the 3D printed sprocket and we had to utilize multiple metal sprockets before the 3D printed sprocket failed.We interpreted the failure of the sprocket in the torsional failure mode as the failure at one cycle of operation. The average torsional failure load across two sprockets was 30.4 ft. lbs.ConfigurationConfiguration 1Configuration 3At torsional failureLoad on the sprocket (W or ft.

The average torsional failure load across two sprockets was 30.4 ft. lbs.ConfigurationConfiguration 1Configuration 3At torsional failureLoad on the sprocket (W or ft. lbs.) 55W32.8W30.4ft.lbs.Load on the sprocket Wt (N)6.5654.54422.032Bending stress (MPa)2.49511.72748.375Number of cycles at failure1.487MM Infinite life (did not experience failure after 3.5MM cycles) 1Table 5: Loads on the sprocket and number of cycles at failure.Note that the bending stress of bulk PLA materials is around 50 MPa. Given these sprockets areat 10% infill, we can estimate the torsional failure bending stress of 100% infill materials to be 3X those f 10% infill sprockets (as supported by the torsional stress experiment results shown in Table 2) or around 25 Mpa. We believe a 50% reduction in the bending stress is reasonable if one were to compare bulk PLA to actual 3D printed PLA due to the magnitude of thermal stresses imposed on the latter.

We believe a 50% reduction in the bending stress is reasonable if one were to compare bulk PLA to actual 3D printed PLA due to the magnitude of thermal stresses imposed on the latter. Given that, our results of observing 8.75 Mpa as the max bending stress is consistent with observations on bending stress in bulk materials. With these data points, a tradeoff curve can be surmised. With loads greater than for Configuration 1, additional data points can be plotted which will be items to explore in a future study.3. ImplicationsThe implications of this study are as follows: - These 3D printed sprockets begin to demonstrate infinite life at loads somewhere between6.5N and 4.5N of load.The results are also consistent with the results of the earlier experiments. The metalsprockets outperform 3D printed sprockets at torsional stress tests while 3D printed sprockets outperform the metal sprockets on fatigue loads.

The metalsprockets outperform 3D printed sprockets at torsional stress tests while 3D printed sprockets outperform the metal sprockets on fatigue loads. It is plausible that there is a load point where the metal and 3D printed sprockets have comparable performance in terms of number of load cycles before they failed, but that could be again be the topic of a future study.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161784G. Future WorkThis study was meant to be just the beginning. We can envision numerous follow-on studies that can both build upon the work we have conducted in this study, and expand upon the study of the impact of 3D printed components.1. Expand fatigue load testsAs discussed in the earlier section, the fatigue load tests can be extended to other load points perhaps with a different power supply and motors and compared against the curves for metal sprockets.2.

Varied Print ParametersThis study can also be extended to other 3D printer build parameters such as build plate temperature, nozzle temperature and cooling time (post printing) to analyze their effects on 3D printed part strength.3. Other MaterialsThis study can also be expanded to other 3D printing materials such as ABS and to othertechnologies such as SLS to quantify the impact they have on performance.4. Testing other types of failure modes and stresses:Though we have tested the two critical failure modes we can extend into numerous other sourcesof failure and stresses such as impact, side shock, offset loads and so on.5. Other StructuresA similar study can be carried out for other part shapes such as gears, axles, hinges, clips, etc.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161785H. AcknowledgementsI would like to acknowledge Mr.

AcknowledgementsI would like to acknowledge Mr. Andy Howell for his tireless guidance with the building of the manufacturing rig, as well as Ms. Rachel Gardner, CEO of Chap Research, and Mr. Eric Rothfus, Chief Mentor of Chap Research, for their valuable reviews and feedback. Last but not the least, I am deeply thankful to Drs. Carolyn Seepersad, Associate Professor of Mechanical Engineering at the University of Texas, Austin, Vikram Devarajan co-founder and CEO of Structured Polymers, Inc. and Jim Mikulak co- founder and VP of Engineering at Structured Polymers, for inspiring me to pursue research in the world of 3D printing.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161786I. References1- Columbus, Louis. "3D Printing Market Forecasts." Forbes. Forbes Magazine, 31 Mar. 2015. Web. 30 Mar. 2016."http://www.instructables.com/id/Simple-Guide-to-Making-Sprockets/2- http://www.firstinspires.org/ 3- Keystone Powdered Metal Company.

Web. 27 Mar. 201616- 3D Printing Blog. "3D Printing Blog." 3D Printing Blog RSS. 3D Printing Blog, 8 Sept. 2014. Web. 27 Mar. 2016. 17- "TETRIX Robotics." TETRIX Robotics. N.p., n.d. Web. 27 Mar. 2016.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161787Appendix A - Sprocket ProductionThe sprockets for these tests were generated by a program called SprocketR. SprocketR is a program created by Chase Probst at Chap Research. The program uses a sine function to generate sprockets. To use this program or learn more about it, visit www.ChapResearch.com.Characterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161788Appendix B - Test Rig for Torsional StressGFigure 12: SolidWorks model of torsional test rig.A – Built from 2x4 wood block, 18 inches long; attached to B B – Built from 2x4 wood block, 3 ft.

long; attached to A, E and D C – 1.125 inch hole that has an Andy-Mark hex bearing in it; a hex axle runs through the bearingwith a sprocket mounted on the axleD – Platform made from wood blocks to provide stability and support through the experiment E – Basket utilized to bear the loads/weights during the experiment F – Metal bar attached to the sprocket 1 ft. from the center of the sprocket; connects the sprocket tothe basket that holds the weightsG – #25 Chain that needs to be kept in tension and fixed to the wood. Our construction included asimple spring mounted to D.Materials List for Torsional Test Analysis 1. 6 feet of 2x4 wood 2. 2 X 2½" hex bore, flanged, heavy duty inner race shielded ball bearing (FR8ZZ-HexHD) (am-2986)3. 3-4 feet of #25 chain 4. 6 inches of ½ inch hex 5. Versa Hub 6. 3 feet of 1x1 aluminum channel 7.

2 X 2½" hex bore, flanged, heavy duty inner race shielded ball bearing (FR8ZZ-HexHD) (am-2986)3. 3-4 feet of #25 chain 4. 6 inches of ½ inch hex 5. Versa Hub 6. 3 feet of 1x1 aluminum channel 7. Scrap metal barCharacterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161789Appendix C – Test Rig for Fatigue Stress AnalysisFigure 13: SolidWorks model of fatigue test rig.A – 18” Aluminum 1x1 channel B – 60 tooth metal sprocket (never changes) C – AndyMark Mini CIM (behind sprocket) D – #25 chain that goes around both sprockets as shown E – 3D printed sprocket that is being tested F – Plate the sprocket is mounted to G – Motor that is back-driven and placed on a loadMaterials List for Fatigue Test Analysis 1. Aluminum 1x1 channel that is 18 inches long 2. 60 tooth metal sprocket, this sprocket never changes 3. AndyMark Mini CIM (Behind sprocket) 4. #25 chain that goes around both sprockets as shown 5. 3D printed sprocket that is being tested 6.

60 tooth metal sprocket, this sprocket never changes 3. AndyMark Mini CIM (Behind sprocket) 4. #25 chain that goes around both sprockets as shown 5. 3D printed sprocket that is being tested 6. Plate the sprocket is mounted to 7. Motor that is back-driven and placed on a loadCharacterization and Testing of 3D Printed SprocketsRohit Srinivasan July 20161790.

ARACTERIZATION OF 2219 ALUMINUM PRODUCED BY ELECTRON BEAM FREEFORM FABRICATIONKaren M. B. Taminger and Robert A. HafleyNASA Langley Research Center, Hampton, VA 23681AbstractResearchers at NASA Langley Research Center are developing a new electron beam freeform fabrication (EB F3) technique to fabricate metal parts. This process introduces metal wire into a molten pool created by a focused electron beam. Potential aerospace applications for this technology include ground-based fabrication of airframe structures and on-orbit construction and repair of space components and structures. Processing windows for reliably producing high quality 2219 aluminum parts using the EB F3 technique are being defined. The effects of translation speed, wire feed rate, and beam power on the resulting microstructures and mechanical properties are explored.

The effects of translation speed, wire feed rate, and beam power on the resulting microstructures and mechanical properties are explored. Tensile properties (ultimate tensile strength, yield strength, and elongation) show little effect over the range of processing conditions tested. Basic processing-microstructure-property correlations are drawn for the EB F3 process.INTRODUCTIONSeveral solid freeform fabrication techniques developed over the past decade have the ability to produce structural metal parts directly from computer-generated designs [1] [2]. These techniques also lend themselves to fabrication of unitized structures that reduce part count, decrease weight, reduce assembly time and cost, and improve structural efficiency, as compared to conventional fabrication and assembly methods. However, many of these techniques rely on lasers, which are ineffective in processing highly reflective metallic materials.

However, many of these techniques rely on lasers, which are ineffective in processing highly reflective metallic materials. Researchers at NASA Langley Research Center have been developing a new electron beam freeform fabrication (EB F3) technique for producing unitized structures from high reflectance aerospace alloys such as aluminum and titanium. The EB F3 process is also conducive for applications in space due to the ease of handling wire feedstock in zero-gravity, high feedstock consumption efficiency, and high power efficiency of this process. Applications in space include on-orbit fabrication of large space structures and repair and manufacturing of spare parts on long duration human exploration missions where resupply can be difficult and costly [3].The EB F3 process uses a focused electron beam to create a molten pool on a metallic substrate. The beam is translated with respect to the surface of the substrate and metal wire is fed into the pool in a layer-additive fashion.

The beam is translated with respect to the surface of the substrate and metal wire is fed into the pool in a layer-additive fashion. The electron beam can be controlled and deflected very precisely and couples very effectively with highly reflective materials. The EB F3 process works with a very finely focused beam that is rastered over a larger pattern to control the size of the molten pool and facilitate capture of the wire. The wire is preheated by the beam, but is not melted until it enters the molten pool. This process is nearly 100% efficient in feedstock consumption and approaches 95% efficiency in power usage.482The EB F3 system at NASA Langley Research Center is comprised of a high power electron beam gun and dual wire feeders capable of independent, simultaneous operation. Positioning is programmable through six axes of motion (X, Y, Z, gun tilt, and positioner tilt and rotate) within a build envelope of 72 in. x 24 in. x 24 in.

Positioning is programmable through six axes of motion (X, Y, Z, gun tilt, and positioner tilt and rotate) within a build envelope of 72 in. x 24 in. x 24 in. The electron beam requires a vacuum on the order of 5x10-5 torr, so the EB F3 system is housed in a large vacuum chamber. This system was used to fabricate 2219 Al deposits using a variety of process conditions, and then the deposits were sectioned to analyze microstructures and tensile properties to correlate with processing parameters.MATERIALSAluminum alloy 2219 is a common aerospace alloy with excellent weldability and good strength and toughness over a wide range of temperatures. Alloy 2319 is the designation for the wire form used as filler wire for welding 2219 Al. The compositions of 2319 Al and 2219 Al are identical (nominally Al-6 wt% Cu) except for slight variations in trace element constituents.

The compositions of 2319 Al and 2219 Al are identical (nominally Al-6 wt% Cu) except for slight variations in trace element constituents. For the purposes of this report, the electron beam deposits will be referred to as 2219 Al because this is the common alloy designation for product forms other than wire and the properties of the EB F3 deposited 2319 Al are directly compared to 2219 Al.EXPERIMENTAL PROCEDURESThe EB F3 process has numerous variables that control the resulting microstructures and shape of the deposit. Three of the most significant of these variables which are easily controlled are the translation speed, wire feed rate, and beam power. These three variables were isolated for detailed microstructural and macrostructural evaluation to understand the effect of processing on the build quality and grain morphology.

These three variables were isolated for detailed microstructural and macrostructural evaluation to understand the effect of processing on the build quality and grain morphology. Deposits were built using seven different combinations of translation speed, wire feed rate, and beam power, then machined for tensile testing.Linear deposits were built from 2319 Al wire onto 0.25 inch thick 2219 Al plate. Deposits for metallurgical analysis and tensile specimens were fabricated ten inches long and one pass wide, with multiple layers to build up to approximately one inch in height. The width of the deposits varied from approximately 0.2 to 0.4 inches, depending upon the processing conditions. The translation direction was selected so that the wire always fed into the leading edge of the molten pool.

The translation direction was selected so that the wire always fed into the leading edge of the molten pool. After each pass the substrate was rotated 180° to maintain the same relationship between the wire, the beam, and the translation vector.A coordinate system related to directions and orientations of the deposit is shown in Figure 1. The longitudinal direction (L) is defined to be parallel to the long axis of travel, the long transverse direction (T) is normal to the surface of the substrate, and the short transverse direction (S) is across the width of the deposited layer. Tensile properties at room temperature were determined using standard 4 in. subsize dogbone specimens per ASTM E8 [4]. Flat tensile specimens were machined in the LT plane and tested in the L direction. The specimen thickness483Figure 1.

subsize dogbone specimens per ASTM E8 [4]. Flat tensile specimens were machined in the LT plane and tested in the L direction. The specimen thickness483Figure 1. Coordinate system related to a typical 2219 Al deposit.was approximately equal to the deposit width, with the surfaces machined flat and parallel to remove surface irregularities introduced by the EB F3 process. Duplicate specimens from each of the seven different processing combinations were tested in the as-deposited condition. In addition, duplicate specimens from four of the seven processing combinations were heat treated to a T62 temper using a standard heat treatment schedule (solutionize at 995°F for one hour, cold water quench, and age at 375°F for 36 hours [5]) and tested.Standard metallographic techniques were employed for microstructural determination. Deposits were sectioned perpendicular to the travel direction and mounted and polished in the ST plane.

Deposits were sectioned perpendicular to the travel direction and mounted and polished in the ST plane. The polished sections were etched with Keller’s reagent to highlight grain morphology. Metallographic specimens were examined using optical microscopy. The chemical composition of the deposits were determined using inductively coupled plasma spectroscopy and compared with the starting 2319 Al wire feedstock.RESULTS AND DISCUSSIONTo evaluate the effect of translation speed, multiple passes were deposited using constant beam power and wire feed rate, while the translation speed was continuously increased from 16 inches per minute (ipm) to 32 ipm during each pass of the deposit. Figure 2a shows the dimensions of this deposit and locations where cross sections were taken for microstructural evaluation. For higher translation speed, the lower heat input and lower volume of wire being fed into the molten pool result in a smaller diameter, shallower molten pool.

For higher translation speed, the lower heat input and lower volume of wire being fed into the molten pool result in a smaller diameter, shallower molten pool. This combined effect produces a decreasing taper in both the width and height of the deposit as the translation speed increases. Figures 2b and 2c show microstructures from the ends of the deposit at the lowest and highest translation speeds, respectively. Light bands of dendrites crossing the grain boundaries define the interpass region, where portions of a previous layer are remelted when a subsequent layer is deposited. A comparison of Figures 2b and 2c reveals that the higher translation speed produces more rapid cooling and results in a homogeneous microstructure with smaller equiaxed grains.

A comparison of Figures 2b and 2c reveals that the higher translation speed produces more rapid cooling and results in a homogeneous microstructure with smaller equiaxed grains. Although dendrites can be seen contained within the grains, pervasive dendrite formation is minimal in the interpass regions of the deposit produced at the higher translation speed.484(a) Photograph of 2219 Al deposit made with increasing translation speed, constant beam power and constant wire feed rate.(b) Section A-A(c) Section B-BFigure 2. Effect of varying translation speed on geometry and microstructure of 2219 Al deposit.To evaluate the effect of wire feed rate, multiple passes were deposited using constant beam power and translation speed, while the wire feed rate was continuously decreased from 120 ipm to 60 ipm during each pass of the deposit. Figure 3a shows the dimensions of this deposit and locations where cross sections were taken for microstructural evaluation.

Figure 3a shows the dimensions of this deposit and locations where cross sections were taken for microstructural evaluation. Since the beam power and translation speed are the same for this case, the energy density is constant across the entire deposit. Increasing the wire feed rate increases the volume of material being fed into the molten pool, thereby increasing the percentage of the energy being used to melt the wire and reducing the amount of energy being absorbed by the part. This results in a narrowing of the molten pool and thus the deposit width, but a significant increase in the deposit height with increasing wire feed rate. Figures 3b and 3c show microstructures from the ends of the deposit at the fastest and slowest wire feed rates, respectively.

Figures 3b and 3c show microstructures from the ends of the deposit at the fastest and slowest wire feed rates, respectively. With the higher wire feed rates, the cooling rate is increased providing a smaller equiaxed grain structure with good homogeneity and no evidence of dendrite growth in the interpass region, as compared to the microstructure developed at the lower wire feed rate. For the lower wire feed rate, the grain size is much larger and less defined.485(a) Photograph of 2219 Al deposit made with decreasing wire feed rate, constant beam power and constant translation speed.(b) Section A-A(c) Section B-BFigure 3. Effect of varying wire feed rate on geometry and microstructure of 2219 Al deposit.The effect of deposition rate was examined by comparing two deposits, each built with constant parameters throughout the build.

The deposit with the lower deposition rate, shown in Figure 4a, was deposited using moderate beam power, moderate wire feed rate, and low translation speed. The deposit with the high deposition rate, shown in Figure 4b, was built using the same translation speed, but 50% higher beam power and 67% higher wire feed rate. The combined effect of higher beam power and higher wire feed rate is an increase in the width and height of the deposit. Increasing the beam power increases the heat input into the part, resulting in a deeper and wider molten pool. The increased wire feed rate adds more material to the molten pool, resulting in the greater deposition rate. Figures 4c and 4d show the microstructural features of the deposits built with lower and higher deposition rates, respectively. The higher deposition rate (Figure 4d) produces a larger grain size with significant dendrite growth, and inhomogeneity in the interpass region.

The higher deposition rate (Figure 4d) produces a larger grain size with significant dendrite growth, and inhomogeneity in the interpass region. There is also evidence of molten material rolling over the edge of previous layers, particularly on the part built at the higher deposition rate. This is a result of the higher heat input increasing the size of the molten pool so that the effect of gravity dominates over surface tension, and the increased fluidity of the aluminum coupled with this gravity effect causes the molten pool to flow over the side of the preceding layer.486(a) Photograph of 2219 Al deposit made with low deposition rate.(b) Photograph of 2219 Al deposit made with high deposition rate.(c) Section A-A(d) Section B-BFigure 4.

Effect of varying deposition rate on geometry and microstructure of 2219 Al deposit.Figure 5 shows the ultimate tensile strength, 0.2% offset yield strength, and total elongation to failure for the EB F3 2219 Al deposits as compared to typical handbook data for sheet and plate products [5]. The data for the as-deposited 2219 Al were averaged over duplicate tests for seven combinations of beam powers, translation speeds, and wire feed rates. Despite the wide range of processing conditions, the majority of the as-deposited 2219 Al data fell in a tight band, as shown by the range bars in Figure 5. The properties of as-deposited 2219 Al fell between those for 2219 Al sheet and plate in the annealed (O temper) and solutionized and naturally aged (T4 temper) tempers. The 2219 Al deposits in the T62 temper also had a very487Figure 5.

The 2219 Al deposits in the T62 temper also had a very487Figure 5. Tensile properties at room temperature of EB F3 deposited 2219 Al as compared to typical handbook values for 2219 Al sheet and plate [5].tight data range, and were equivalent to typical T62 handbook properties for sheet and plate product [5].Over the entire range of beam powers, translation speeds, and wire feed rates tested, the deposits exhibited very little porosity, no cracking, and complete fusion between deposited layers. The chemical composition of the deposits compared with the starting 2319 Al wire feedstock confirmed that the alloy compositions were unaffected by the EB F3 process. The tensile properties were also very consistent over a wide range of processing conditions. These results indicate that the EB F3 process is robust and provides a wide processing envelope from which 100% dense deposits can be made with useful mechanical properties.

These results indicate that the EB F3 process is robust and provides a wide processing envelope from which 100% dense deposits can be made with useful mechanical properties. However, several processing combinations resulted in excess heat input into the molten pool, as evidenced by the widening of the molten pool, roll over at the edges of previous deposited layers, and dendrite formation and growth. Further process development is planned to optimize the thermal input to reduce distortion and residual stresses, and to improve the microstructure and mechanical properties of the resulting EB F3 deposited materials.CONCLUSIONS1. The potential for building 2219 Al structures using the EB F3 process was successfully demonstrated.2. The effects from varying the EB F3 processing conditions were correlated with the resulting microstructures and deposition features of 2219 Al deposits.4883.

The effects from varying the EB F3 processing conditions were correlated with the resulting microstructures and deposition features of 2219 Al deposits.4883. Tensile properties of the 2219 Al deposits produced using EB F3 showed little variation over the wide range of processing conditions tested.4. After T62 heat treatment, the 2219 Al deposits produced using EB F3 possessed tensile properties comparable to typical handbook data for T62 sheet and plate product.5. The EB F3 process is robust, as evidenced by the wide range of processing conditions to achieve defect-free deposits and good tensile properties.6. Process optimization is required to improve thermal management in the 2219 Al EB F3 deposits.REFERENCES[1] J. W. Sears, “Direct Laser Powder Deposition – State of the Art,” Powder Materials: Current Research and Industrial Practices, Proceedings of the 1999 Fall TMS Meeting, Ed. By F. D. S. Marquis, November 1999, pp. 213-226.[2] P. F.

ARACTERIZATION OF ACOUSTIC SOFTENING OF ALUMINUM 6061 WITHIN A PLASTICITY FRAMEWORKQ. Mao, N. Coutris, and G. Fadel Department of Mechanical Engineering, Clemson University, SC 29630 REVIEWEDAbstractUltrasonic additive manufacturing (UAM) is a rapid prototyping technology that features a metal joining process through ultrasonic welding. The bonding mechanisms and mechanics of UAM have been investigated for decades. Meanwhile, the plastic deformations of metals were extensively studied by many researchers for their significant roles in bond formation. However, most of these research efforts considered solely the surface frictional effects on plastic deformation whereas the volumetric effects of ultrasound were rarely considered. This paper investigates the effects of ultrasound on deformation of Aluminum 6061 through experimental studies and highlights the volumetric effects of ultrasound, i.e.

This paper investigates the effects of ultrasound on deformation of Aluminum 6061 through experimental studies and highlights the volumetric effects of ultrasound, i.e. the “acoustic softening”: a stress reduction on the stress-strain relation of Aluminum 6061 upon application of ultrasounds. Based on observations obtained from a designed experimental setup, a phenomenological model is proposed to characterize the acoustic softening effects in terms of the ultrasonic intensities. Additionally, by modifying Hockett’s plasticity model, a plasticity frame work is established to characterize the deformation of Aluminum 6061 in UAM. The acoustic softening model is then incorporated into the plasticity framework. The complete model is then validated by comparing its predictions with experimental measurements.1.

The acoustic softening model is then incorporated into the plasticity framework. The complete model is then validated by comparing its predictions with experimental measurements.1. IntroductionThe ultrasonic additive manufacturing (UAM) is a solid state fabrication process that combines an additive process of joining metal foils through ultrasonic metal welding and a subtractive process of CNC contour milling. The additive process bonds thin metal foils of 100 – 150 µm thickness to the substrate layer by layer with combined application of ultrasonic energy and compression, whereas the subtractive process shapes the deposited layers to the required contour with CNC milling. The bonding mechanism of UAM has been studied for decades yet no uniform conclusion has been achieved. Metallurgical adhesion is supported by many researchers [1-7] as the bonding mechanism during which material plastic deformation plays a vital role in closing and filling voids at the contact interface.

Diffusion across the weld interface is supported by some investigations based on the observation of high plastic strain rate during the ultrasonic welding near the bonding interface, which significantly enhances diffusion by increasing vacancy concentrations within materials [8, 9, 10]. Similar observations are made by other researchers, who claim that the high vacancy concentration caused by the high strain rate significantly depresses material melting temperature, thus allowing localized melting to occur [10]. Recrystallization is also proposed as a cause of bonding [5, 10, 11]. It is believed that plastic deformation and temperature rise due to the continuous input of ultrasonic energy provide the necessary driving force for recrystallization. Mechanical interlocking is reported by a few researchers [7] who studied the bonding of dissimilar materials in a combination of one material437being soft and the other hard. Severe plastic deformation is observed in the soft material.

Severe plastic deformation is observed in the soft material. Despite the divergence among the competing theories, all the investigations show the significance of plastic deformation of materials in promoting bonds formation regardless of its causes. It is widely accepted that plastic deformation promotes bonds formation by dispersing surface oxides and contaminants, increasing contact areas of pure metal, and maintaining the already formed bonds [2, 5]. As a result, the material plasticity in UAM requires a comprehensive investigation and characterization to fully understand the bonding process.Ultrasonic energy, being the major energy input of UAM, affects material plasticity in terms of surface and volumetric effects. The surface effect includes friction-induced plastic deformation at the mating surface and frictional heating that reduces the plastic yield stress of materials by increasing temperature, i.e. the “thermal softening”.

the “thermal softening”. The volumetric effect refers to a unique effect of ultrasound on the volume of metals: the stress required for plastic deformation is reduced upon application of ultrasound, known as the “acoustic softening”. This effect is believed to affect the plasticity of metals more effectively than thermal softening with less energy input [2, 5, 12-14]. While the surface effects of ultrasound have been extensively studied, the volumetric effects are considered by only a few researchers in studies of UAM [13, 14]. This paper investigates the acoustic softening of aluminum 6061 which is a material extensively used in UAM, and characterizes the acoustic softening using a selected plasticity model, aiming at establishing an analytical material plasticity framework for future studies of UAM.Despite the limited investigations on volumetric effects of UAM, the physics of acoustic softening in metals have been under investigation since the 1950s [15].

The acoustic softening is first documented by Blaha and Langenecker, who applied tensile tests with superposition of ultrasound on a variety of metals: aluminum single crystals and polycrystals, coppers, nickels, magnesium and titanium [12, 16]. According to these researchers, the acoustic softening take places instantaneously as ultrasound starts and the tensile stress drops abruptly. The stress returns to its original value as soon as the ultrasound is switched off. The stress reduction is observed to be “proportional” to the ultrasound intensity [12]. However, when the ultrasonic intensity exceeds a certain threshold, a “residual hardening” is observed, i.e., the stress returns to a value higher than its original value as the ultrasound is switched off.

Additionally, based on the stress- strain examination, the curve is observed to be similar to thermal softening, yet the consumption of ultrasonic energy is of the order of (cid:883)(cid:882)(cid:2874) less than thermal energy for the same amount of stress reduction. Later, many researchers investigated the acoustic softening. Some conducted experiments similar to Langenecker’s but reported different observations [17-20]. For instance, Nevill and Brontzen [17] claimed that stress reduction is a linear function of vibration amplitude rather than being proportional to ultrasonic energy. Others designed different experimental setups and obtained conflicting observations [21, 22]. For instance, Culp conducted compression tests on aluminum alloy 6063 with superposition of ultrasound and reported “residual softening” instead of “residual hardening” [21]. Yao et al.

For instance, Culp conducted compression tests on aluminum alloy 6063 with superposition of ultrasound and reported “residual softening” instead of “residual hardening” [21]. Yao et al. performed compression tests superimposed with sound waves of 9.8 kHz which is below the frequency range of ultrasound and reported residual hardening when the sound wave stopped [22]. The divergences among experimental observations in the literature renders the modeling of acoustic softening challenging and thus experiments on acoustic softening are necessary and crucial in quantifying the stress reduction in terms of ultrasonic energy levels.438In a tensile/compression test, acoustic softening couples with strain hardening in the plastic deformation region. Consequently, the acoustic softening model has to be incorporated into a plasticity model that accounts for strain hardening. Kelly et al.

Consequently, the acoustic softening model has to be incorporated into a plasticity model that accounts for strain hardening. Kelly et al. introduced acoustic softening into a power law model proposed by Hockett which characterizes strain hardening and thermal softening [14]. The power law equations are derived from true stress versus true strain relations of aluminum 1100-O in the plastic deformation region over a range of strain rates and temperatures by varying the coefficient and the exponent [24]. By assuming that acoustic softening linearly affects the true stress, Kelly et al. modified the plasticity model and introduced acoustic softening factor [14]. Though the modified model gives a prediction that resembles the experimental observations by Blaha and Langenecker, the assumption of acoustic softening linearly affecting stress lacks experimental support.

Additionally, the material used in Langenecker’s experiment is aluminum single crystals whereas the material being characterized by Kelly et al. is commercially pure aluminum 1100-O, which has a polycrystalline structure. Siddiq et al. established a thermal-mechanical material model of UAM which is solved using finite element methods [25]. The model characterizes plasticity of aluminum 6061 under cyclic loading. It combines a nonlinear isotropic model and a kinematic hardening model which are proposed by Chaboche and his coworker [26]. The model is readily available in ABAQUS and is subject to modification. The acoustic softening is introduced by modifying the isotropic and kinematic hardening rules. The assumption made about acoustic softening is that the stress reduction is proportional to ultrasonic intensity, which is reported by Blaha and Langenecker, and Green. [16, 27]. However, the model is subject to debate since Siddiq et al.

[16, 27]. However, the model is subject to debate since Siddiq et al. model for aluminum 6061 using experimental data of aluminum single crystals [25]. By comparing the experiment data and those reported by Blaha and Langenecker, this paper demonstrates that the stress reduction in polycrystalline aluminum alloy is significantly different from that in aluminum single crystals. Other models that have been applied to UAM include crystal plasticity models [22, 28]. By using the Crystal Plasticity Finite Element Method (CP-FEM) which is readily available in ABAQUS, a cube of polycrystalline material is constructed using the electron backscattered diffraction (EBSD) data of aluminum alloy 6061[28]. The acoustic softening is introduced by modifying the plastic flow rule and by adding an ultrasonic softening term which is a function of ultrasound intensity.

The acoustic softening is introduced by modifying the plastic flow rule and by adding an ultrasonic softening term which is a function of ultrasound intensity. The assumption about the softening term comes from the observation by Blaha and Langenecker that the decrease in yield limit is directly proportional to the ultrasound energy input. However, the softening term is derived as a polynomial of ultrasound intensity. The softening term is not proportional to ultrasound intensity and it is difficult to determine the order of the polynomial without experimental data. Yao et al. utilized the thermal activation model proposed by Kocks and the dislocation evolution law proposed by Krausz and Krausz to account for acoustic softening and residual hardening within a crystal plasticity frame work [22, 29, 30].

The underlying assumptions are that the dislocations overcome obstacles with the assistance of thermal energy as ultrasound starts and that the dislocations evolve and tangle with each other to cause residual hardening as ultrasound stops [22]. The acoustic softening model relies heavily on the assumed mechanism of thermal activation at dislocations whereas the actual mechanisms of acoustic softening remain subjective to debate.In summary, it is shown that the bonding process of UAM is significantly influenced by plasticity of the materials in use and that the material plasticity is influenced by surface and volumetric effects of ultrasound. However, the effect of ultrasound on UAM, specifically the439volumetric effect has not been fully investigated and characterized. The existing literature reports diverging observations of acoustic softening on different metallic materials, making it difficult to characterize for a specific metal without conducting experiments.

The existing literature reports diverging observations of acoustic softening on different metallic materials, making it difficult to characterize for a specific metal without conducting experiments. Most of the existing analytical models of acoustic softening reviewed lack support of experimental data. To the best knowledge of the authors, no experimental investigations have been reported regarding acoustic softening of aluminum 6061. In this paper, an experimental setup similar to Langenecker’s is used to observe acoustic softening on aluminum 6061. Stress reductions in plastic region are measured and stress strain curves are recorded at different levels of ultrasound intensities. Based on the experimental observations, an analytical phenomenological model is proposed to characterize the acoustic softening in the plastic region combined with strain hardening effect.

Based on the experimental observations, an analytical phenomenological model is proposed to characterize the acoustic softening in the plastic region combined with strain hardening effect. By comparing the predictions from the analytical model and the experimental data, it is shown that the proposed model is able to characterize the acoustic softening effect with good accuracy.2. Experimental SetupFig.1 shows the scheme of the experimentation which is similar to Langenecker’s testing apparatus. It consists of an MTS hydraulic tensile test machine, a Branson 2000 ultrasonic plastic welder and a rigid frame made of aluminum 6061. The MTS tensile test machine applies tensile or compression force in the vertical direction to the bottom end of the specimen while the ultrasonic plastic welder provides longitudinal ultrasonic vibration to the top of the specimen. A load cell that connects the MTS actuator and the specimen is used to record the loading profile.

A load cell that connects the MTS actuator and the specimen is used to record the loading profile. The ultrasonic plastic welder consists of an ultrasonic transducer that converts electric energy into ultrasonic vibration with a fixed frequency of 20 kHz, a booster that amplifies vibration and an ultrasonic horn that further amplifies vibration and delivers ultrasonic energy to the specimen. A circular high gain horn made of aluminum heat-treated alloy is used and the maximum vibration amplitude is 48 microns. Due to the fact that the ultrasonic welder is designed to take compressional load rather than tensile load, the ultrasonic horn cannot be connected to the specimen that is tensioned by the MTS machine. As a result, a box frame made of aluminum 6061 plates with thickness of 0.5 inches is designed to provide a mounting position for the specimen and to take the tensile load off the ultrasonic welder.

As a result, a box frame made of aluminum 6061 plates with thickness of 0.5 inches is designed to provide a mounting position for the specimen and to take the tensile load off the ultrasonic welder. The specimen is installed such that it is oscillated by the ultrasonic horn at the top end, fixed at a point right below the oscillating point and is tensioned at the bottom end. The specimen is in cylindrical shape with two fillets to transit from large installation diameter to small testing diameter without causing undesired flexure in the machining process (fig.2). In the gage section, the diameter is 3/16 inch and the length is 1 inch. The specimen is made of aluminum 6061 for its extensive application in UAM. Apart from the testing system, a Polytec laser vibrometer is used to measure the maximum vibration amplitude available in the test.

2 Specimen design (a) specimen dimensions (unit: inch), (b) the actual specimen3. Experimentation Details and ObservationsIn order to characterize the acoustic softening in the plastic deformation region under different levels of ultrasonic energy in aluminum 6061, a series of tensile tests are run as the input of ultrasonic energy varies. The specimen is subjected to static tensile load at a constant loading speed of 0.1 mm/s until it fails. Meanwhile, the ultrasonic welder is started when the specimen begins to yield. As the welder starts, the ultrasonic horn is brought into contact with the upper end of the specimen at a constant speed and is compressed on top of the specimen with a moderate force of around 100 lbs. Then the ultrasound starts and the longitudinal ultrasonic wave propagates through the specimen that is being tensioned. The ultrasonic irradiation lasts for 10 seconds before it stops. The compression is then removed and the ultrasonic horn is brought up.

The ultrasonic irradiation lasts for 10 seconds before it stops. The compression is then removed and the ultrasonic horn is brought up. The amount of ultrasonic energy applied during the irradiation is recorded by the welder controller and can be retrieved when the test completes. Among the 26 tests that are completed with success, 6 tests are selected to present their stress strain curves at different levels of ultrasonic energy input: 0 J, 915J, 2684 J, 3718 J, 5506 J and 6040 J. By setting the maximum ultrasonic energy input as the 100% energy level, the energy inputs are converted into percentages for clarity. Fig. 3 shows the engineering stress strain curves of aluminum 6061 at 6 different levels of energy input. As the ultrasonic energy level increases, the stress reduction increases accordingly and the material becomes “softer”. In the four lowest energy levels, the softening curves show a combined effect of acoustic softening, i.e. stress reduction, and strain hardening.

In the four lowest energy levels, the softening curves show a combined effect of acoustic softening, i.e. stress reduction, and strain hardening. In the other two high energy levels however, the strain hardening diminishes and the441curves become straight lines, which are similar to curves reported by Blaha and Langenecker when ultrasonic intensity is large [12]. During the experiment, a thermal camera is used to record the temperature gradient at the acoustic softening spot. Two thermal couples are glued to the surface of the spot to identify the temperature increase. The room temperature is 70 (cid:1324) and the highest temperature observed during ultrasonic irradiation is 120(cid:1324). Due to the high temperature (400 (cid:1324)) required for thermal softening to show significance [16], the thermal softening in our experiment is considered as insignificant.Fig.4 shows the comparison of our experimental data with that of Blaha and Langenecker.

2) The ultrasound intensities used in our experiment are two orders of magnitude higher than the ultrasound intensity used in Langenecker’s experiment [12], yet the stress reductions achieved in our experiment are much lower. 3) Instead of residual hardening, moderate residual softening is observed in stress strain relations as soon as the ultrasound stops. The residual softening then dies down as plastic deformation continues, possibly due to the dissipation of previously stored acoustic energy [21]. These observations demonstrate that the acoustic softening in aluminum alloys is quantitatively different from the acoustic softening in aluminum single crystals. It takes much higher ultrasonic energy to achieve the same stress reduction in aluminum 6061 than in aluminum single crystals.

(a) Our experimental data (b) Langenecker’s experimental data4. Analytical ModelWith the acoustic softening effect observed and experimental data collected, a phenomenological model is required to incorporate acoustic softening effect into the stress strain relations of the material used for studies of UAM. The stress strain curves present deformation in elastic and plastic regions. In the elastic deformation region, no ultrasound is applied in this experiment since elastic deformation is not the focus of this paper. According to previous studies [23] however, the application of ultrasound does not change the Young’s modulus of material. Therefore the value of the Young’s modulus is calculated from the true stress strain data E = 95646 ksi. In the plastic deformation region, acoustic softening is observed to be coupled with strain hardening. As a result, the analytical model has to incorporate an equation that characterizes the hardening in the stress and strain relation.

As a result, the analytical model has to incorporate an equation that characterizes the hardening in the stress and strain relation. Hockett utilized power law equations to characterize the relation between stress and strain as strain rate and temperature change [24]. The equation is shown to be able to effectively capture the hardening curve measured in experiments. Therefore we introduce the power law equation to account for the hardening effect in our experiment:(cid:3041) (2) (cid:2026)(cid:3047) (cid:3404) (cid:1837)(cid:2013)(cid:3047) Where (cid:2026)(cid:3047)(cid:3)the instantaneous true stress is required for plastic deformation, (cid:2013)(cid:3047) is true strain, (cid:1837) is the strength coefficient and (cid:1866) is the strain hardening coefficient. (cid:1837) and (cid:1866) are constants determined by fitting the power law equation to the stress-strain data free of ultrasonic irradiation.

The true stress versus true strain curves are shown in fig.5. The power law equation is fitted to the stress strain data without ultrasonic irradiation using least square regression. The regression gives the strain hardening coefficient(cid:3)(cid:1866) (cid:3404) (cid:882)(cid:484)(cid:882)(cid:888). With the hardening model ready, the softening model is investigated by expressing the stress reduction in terms of the ultrasonic energy input. The relation betweenthe strength coefficient (cid:1837) (cid:3404) (cid:888)(cid:484)(cid:890)(cid:889) (cid:3400) (cid:883)(cid:882)(cid:2872) and443stress reduction and ultrasonic energy input is first observed by Blaha and Langenecker [12]. Later, several researchers investigated the relation and proposed different models [13, 21, 22]. Moreover, it is reported by Blaha and Langenecker that the stress reduction is observed to be “directly proportional” to ultrasonic energy input [12].

Moreover, it is reported by Blaha and Langenecker that the stress reduction is observed to be “directly proportional” to ultrasonic energy input [12]. Therefore we start investigating the relation between stress reduction and energy input with linear regression.60000True stress-true strain curveUS = 0 J (0%)5000040000) i s p (30000s s e r t S2000010000US = 915 J (15.1%) US = 2684 J (44.4%) US = 3718 J (61.6%) US = 5506 J (91.6%) US = 6040 J (100%)000.010.020.030.040.05StrainFig.5 the true stress- true strain curvesFig. 6 shows the relation between stress reduction and energy input. The horizontal axis represents the ultrasonic energy input (cid:8)(cid:2931) and the vertical axis represents the stress reduction ratio(cid:3)(cid:580).

However, within the strain range in test, the change of stress due to strain is much smaller than the stress reduction due to acoustic softening. As a result, the stress reductions are approximated as constant values which are calculated by averaging the stress reduction at true strain(cid:3)(cid:578)(cid:2930)(cid:2928) (cid:3404) (cid:882)(cid:484)(cid:882)(cid:884)(cid:481) (cid:882)(cid:484)(cid:882)(cid:885). The calculations are shown in Table 1. As ultrasonic energy increases, the softening curves tend to lose the hardening effect and transit to straight lines with slope. As a result, only curves with the five lowest ultrasonic energy levels are investigated. The other curve is modeled separately.

The true stress strain curves are truncated and only the part of curves with445acoustic softening are displayed. The solid lines represent experimental data and the lines with markers represent the approximation made by the analytical models. It can be observed that the prediction from the analytic models agrees well with experimental data in the four lowest energy levels. In the fifth energy level, (cid:8)(cid:2931)=5560 J, the strain hardening is not significant at the beginning and the curve levels off. As deformation proceeds, the power-law shape strain hardening appears and the curve starts to agree with the prediction from the analytical model. In the last and the highest energy level, however, the acoustic softening is so significant that the power-law shape of the strain hardening is transformed into the linear shape of hardening. As a result, a linear hardening model has to be introduced instead for characterization.

As a result, a linear hardening model has to be introduced instead for characterization. A good agreement is shown between the linear model prediction and the experimental data.In summary, an analytical material model of aluminum 6061 is proposed to account for both acoustic softening and strain hardening that occur simultaneously as material deforms in plastic region. The ultrasound-induced stress reduction is proportional to ultrasonic energy input. The strain hardening behavior of the material, however, differs depending on the values of ultrasonic energy. With ultrasonic energy below a certain threshold, in this case 60% of the maximum energy input, the strain hardening displays a curved shape and can be modeled using power law equations.

With ultrasonic energy below a certain threshold, in this case 60% of the maximum energy input, the strain hardening displays a curved shape and can be modeled using power law equations. With ultrasonic energy above the threshold of 60%, however, the strain hardening diminishes and has to be modeled using a linear hardening model.5.65.55.4x 104Experimental E = 0 J Experimental E = 916 J Experimental E = 2684 J Experimental E = 3718 J Experimental E = 5560 J Experimental E = 6040 J Analytic E = 0 J Analytic E = 916 J Analytic E = 2684 J Analytic E = 3718 J Analytic E = 5560 J Analytic E = 6040 J5.3) is p ( s s e r t5.2S e u r T5.154.94.80.010.0150.02True Strain0.0250.03Fig. 7 The comparison of experimental data and prediction by analytical model5. DiscussionIn this paper, a series of experiments are conducted to characterize the acoustic softening of aluminum 6061 under different levels of ultrasonic energy input.

DiscussionIn this paper, a series of experiments are conducted to characterize the acoustic softening of aluminum 6061 under different levels of ultrasonic energy input. However, some advantages and limitations of the experimentation need to be pointed out. First, the experimentation setup used in this paper proves to be more effective in reducing noise from thermal softening than4460.035other designs [23]. The experimental setups for studies of acoustic softening vary in the literature. However, the method to observe acoustic softening is based on the stress strain diagram of materials. As a result, tensile or compression tests are conducted to achieve stress strain relations while transversal or longitudinal ultrasonic waves are applied to show the significance of ultrasound. Blaha and Langenecker, and some researchers [12, 27, 31] performed tensile tests on specimens while applying longitudinal ultrasonic waves.

Blaha and Langenecker, and some researchers [12, 27, 31] performed tensile tests on specimens while applying longitudinal ultrasonic waves. Culp and other researchers [21, 31-33] replaced tensile tests with compression tests while applying longitudinal ultrasonic waves. Siu et al. and Yao et al.[22, 34] proposed setups that apply ultrasonic wave in transversal directions during compression tests. Dutta et al. applied longitudinal ultrasonic waves in a direction perpendicular to the axis of the specimen during tensile tests. By evaluating these experimental setups, it is found that compression tests introduce more friction than tensile tests, and that transversal ultrasonic vibrations introduce more friction than longitudinal vibrations. As a result, the combination of tensile tests and application of longitudinal vibration reduces the noise from thermal softening by avoiding unnecessary frictions.

As a result, the combination of tensile tests and application of longitudinal vibration reduces the noise from thermal softening by avoiding unnecessary frictions. Second, the ultrasonic energy levels used in the analytical model are measured from the electric power consumption which does not indicate the exact amount of energy that softens the specimen. It is extremely difficult to quantify the exact amount of energy input into specimen due to the difficulties in quantifying the energy losses during ultrasound propagation in the ultrasonic stack (transducer, booster and horn), specimens and frames. However, the objective of the analytical model is to characterize acoustic softening by means of investigating the relation between the stress reductions and the energy absorptions. By replacing the ultrasonic energy input with electric energy consumption, it is justifiable to assume that the relation between stress reduction and energy input remain unchanged.

By replacing the ultrasonic energy input with electric energy consumption, it is justifiable to assume that the relation between stress reduction and energy input remain unchanged. Therefore the electric energy input serves only as an indicator of ultrasonic energy input for the purpose of characterizing acoustic softening. Third, the geometry of the tensile specimens in use are different from that of ASTM standard specimens due to manufacturing limitations. Two fillets are designed for the specimen to transit large installation diameter to small testing diameter.According to experimental observations, the stress is reduced by as much as 10%, which is not as significant as the reductions reported in the literature. Blaha and Langenecker observed stress reduction as high as 100% in aluminum single crystals [12]. Culp obtained approximately 50% stress reduction in aluminum 6063, 30% stress reduction in steel 1018 and less than 10% in magnesium [21].

Culp obtained approximately 50% stress reduction in aluminum 6063, 30% stress reduction in steel 1018 and less than 10% in magnesium [21]. Daud achieved approximately 25% stress reduction in aluminum 1050 [31]. Yao obtained roughly 50% stress reduction in aluminum 1100 [22]. In comparison, the aluminum 6061 has much higher yield strength than the materials in which large stress reductions are achieved and thus requires more energy input to achieve similar stress reduction. However, since the ultrasonic energy input is not available in many of the studies, the quantitative comparison becomes difficult.6. Conclusion and Future WorkBy designing the experimental setup that superimposes longitudinal ultrasonic wave to tensile test, the stress reduction of aluminum 6061 is measured at different levels of ultrasonic energy input.

It is observed that:4471) The acoustic softening in aluminum 6061 is small compared with those in aluminum single crystals whereas the applied ultrasonic energy is much higher than that applied to aluminum single crystals. 2) A residual effect of softening is observed as soon as the ultrasound stops. The effect then diminishes as material deformation continues. 3) The stress reduction is observed to be linearly proportional to the ultrasonic energy input. Based on the observation, an analytical model is established to characterize the acoustic softening effect. The power law model is able to capture the strain hardening as ultrasonic energy input is below a certain threshold; a linear hardening model is introduced to capture the strain hardening when ultrasonic energy input is high.

The stress reduction is modeled by using a “softening term” which is proportional to ultrasonic energy input.With the acoustic softening experimentally observed and analytically characterized, the future work includes observing and characterizing thermal softening effects, and establishing a UAM model in which both acoustic softening and thermal softening are implemented.7. AcknowledgementThis research was partially supported by the National Science Foundation under grant # 1068977 entitled “Understanding the Macroscopic Dynamics of Ultrasonic Consolidation”. The views presented in this work do not necessarily reflect those of our sponsor whose support is gratefully acknowledged. We also wish to thank Ms. Sophie Morneau from Branson Ultrasonics Corporation for loaning some ultrasonic component.References[1] Kong, C., Soar, R., & Dickens, P. (2003). Characterisation of aluminium alloy 6061 for the ultrasonic consolidation process.

aracterization of Bulk to Thin Wall Mechanical Response Transition in Powder Bed AMBen Brown1, Wes Everhart1, and Joe Dinardo11Department of Energy’s National Security Campus Managed by Honeywell FM&T, Kansas City, MO 64147Notice: This manuscript has been authored by Honeywell Federal Manufacturing & Technologies under Contract No.DE-NA-0000622 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid-up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposesAbstractIn the development of powder bed AM process parameters, the characterization of mechanical properties is generally performed through relatively large mechanical test samples that represent a bulk response.

This provides an accurate representation of mechanical properties for equivalently sized or larger parts. However as feature size is reduced, mechanical properties transition from a standard bulk response to a thin wall response where lower power border scans and surface roughness have a larger eﬀect. This study identiﬁes this threshold between bulk and thin wall for 304L SS on the Selective Laser Melting (SLM) platform and Ti-6Al-4V on the Electron Beam Melting (EBM) plat- form. A possible method for improving those properties and shifting the transition from bulk to thin wall response to smaller wall thicknesses was investigated. Mechanical test- ing and fractography was performed on samples to characterize the eﬀect of wall thickness.IntroductionWith the recent advancements in powder bed additive manufacturing (AM) equip- ment, previously unmanufacturable geometries are now a possibility.

Often, the designs most conducive for economically viable AM parts have ﬁne features such as lattice struc- tures or thin walls. The production of these geometries represents a unique problem in the determination of mechanical properties. In this study samples produced by Selective Laser Melting (SLM) and Electron Beam Melting (EBM) were used to investigate the501eﬀect of reducing wall thickness on resultant tensile properties.The SLM process uses a laser in an inert atmosphere to selectively melt layers of loose metal powder into a solid. EBM also fabricates parts from a bed of loose metal pow- der, however the energy source is an electron beam necessitating a vacuum environment. Both of these technologies use scan patterns that include unique hatch ﬁll and border scan parameters.

Both of these technologies use scan patterns that include unique hatch ﬁll and border scan parameters. Typical hatch ﬁll parameters are developed to achieve dense parts at high build rates and border parameters are developed to smooth the outer surface of the hatch ﬁll scan area resulting in better surface ﬁnish. Because of the function of the border scans, they will not necessarily have a suﬃcient parameter combination to provide adequate interlayer adhesion on their own, resulting in poor mechanical properties. As a part’s thickness reduces, an increasingly greater percentage of the cross sectional area is built with the non-structural border parameters.Though the eﬀect of ﬁne feature geometry may signiﬁcantly impact part performance, very little work has be done in this area for AM produced parts.

Examples of the large bulk specimen used in many standard development practices include use of 10 mm x 10 10 mm tensile specimen used by Niendorf et mm x 7 mm density pillars [4] and the al. [6]. Ahujaa et al. [3] demonstrated an ability to build walls at 100 µm thick and used these machine parameters to build larger bulk samples. This was accomplished with only single line hatch ﬁll scans and no smoothing contours, resulting in an apparent rough surface ﬁnish. Ideally for end use ﬁne feature parts, fully dense and smooth geometries would be produced. Because AM produces a relatively rough surface ﬁnish under ideal conditions, it is especially prone to decreased mechanical properties and surface initiated failure. Krauss et al. [5] demonstrated that thin walls can exhibit a very rough surface if not compensated for and that the manufacturability of thin walls is orientation depen- dent. Song et al.

Krauss et al. [5] demonstrated that thin walls can exhibit a very rough surface if not compensated for and that the manufacturability of thin walls is orientation depen- dent. Song et al. [9] optimized processing parameters to achieve a desired wall thickness based on input energy. Qui et al. [8] produced thin lattice structures that were then mechanically tested in compression. This study demonstrated the relationship of laser parameters on resulting part porosity and geometry. Abele et al. showed how increased porosity leads to a reduction in mechanical properties in large samples and at thin geome- tries porosity creates a threshold for achieving gastight parts [2]. For comparison to thin wall tensile testing of similar materials with diﬀerent processing methods, the work of Pardoen et al. [7] showed that the mechanical response of 316L is not sensitive to change in thickness down to 0.63 mm.

[7] showed that the mechanical response of 316L is not sensitive to change in thickness down to 0.63 mm. From these previous studies it can be seen that using a standard set of parameters results in a threshold for mechanical properties when reducing wall thickness. By identifying and characterizing the reasons for this threshold where the response changes, scan patterns can be manipulated to reduce or even eliminate this eﬀect.(cid:31)Experimental SetupFor this research, a Renishaw AM250 SLM system and an Arcam A2X EBM system were used. The AM250 uses a 1070 nm ﬁber laser with a maximum output of 200 W and an approximate spot diameter of 70 µm. The 304L powder used was 15-45 µm diameter. The Renishaw AM250 uses a point exposure scan pattern where a single point is exposed, the laser is turned oﬀ, repositioned, and then the next point is exposed. The next layer502is then exposed in the same way, however the pattern is rotated a speciﬁed amount.

The next layer502is then exposed in the same way, however the pattern is rotated a speciﬁed amount. A diagram of this Standard scan pattern is illustrated in Figure 1.Figure 1: Schematic of the point wise AM250 Hatch Fill Scan PatternFigure 2: Schematic of the AM250 Border Scan Oﬀsets for one layer the Standard Parameter SetIn addition to the illustrated hatch ﬁll pattern, two border scans are used to smooth the outer proﬁle. A schematic of hatch ﬁll and border scans can be found in Figure 2. A spot compensation is also applied to the outer most border scan. This distance compensates for the width of the laser spot, allowing for a more accurate geometry on the part. With this scan pattern, laser power, point distance, exposure time, and oﬀsets between subsequent hatch ﬁll and border scans are all user deﬁned. Previously developed 304L parameters sets were used for building samples on the AM250.

Previously developed 304L parameters sets were used for building samples on the AM250. A summary of the diﬀerences between hatch ﬁll and border scan parameters can be found in Table 1. The hatch spacing between hatch ﬁll scans was set to 85 µm, border distance was 40 µm, and the ﬁll hatch oﬀset was 45 µm.Table 1: AM250 Scan ParametersMachine Parameter Hatch Fill BorderLaser Power (watts) Point Distance (µm) Exposure Time (µsec)200 85 53150 75 100503The Arcam A2X uses a 3000 W electron beam output with an approximate spot size of 200 µm. The vacuum chamber is pumped down to 10−4 mbar pressure when build- ing. The Ti-6Al-4V samples built on the A2X used stock parameters and powder 45-100 µm diameter. The scan pattern used by the A2X is an orthogonal hatch for the border followed by a hatch ﬁll scan.

The Ti-6Al-4V samples built on the A2X used stock parameters and powder 45-100 µm diameter. The scan pattern used by the A2X is an orthogonal hatch for the border followed by a hatch ﬁll scan. While border scans and hatch ﬁll scan utilize separate parameter sets, speciﬁc comparison for current and speed are not discussed due to the complexity of parameters and their dependence on part geometry.Flat samples conforming to ASTM E8 [1] with a gage length of 50.0 mm and gage width of 12.5 mm were produced in the vertical orientation at the varying model thick- nesses of 4.0, 3.0, 2.0, 1.0, 0.75, 0.5, and 0.25 mm. Samples were oriented perpendicular relative to the recoating direction of each machine, resulting in a greater stiﬀness. Each sample was produced in triplicate, removed from the build plate, and mechanically tested per ASTM E8. Due to the rough surface of the samples, the cross sectional area exhibits error.

Each sample was produced in triplicate, removed from the build plate, and mechanically tested per ASTM E8. Due to the rough surface of the samples, the cross sectional area exhibits error. Scanning electron microscopy was performed on fracture surfaces and sample cross sections.Results and DiscussionThe ﬁrst set of SLM 304L mechanical testing can be found in Figure 3. It should be noted that samples were overbuilt due to lack of spot compensation from the mod- eled geometries previously listed and that all strength calculations were performed using measured thicknesses.

Thickness on Figures 3, 12, and 10 all list the average measured thickness by digital calipers from each set of three samples.100120100Yield Yield UTS UTS Elongation80) I S K (h t g n e r t S8060406040E l o n g a t i o n(%)2020003.94 1.08 0.35 Average Measured Sample Thickness (mm)0.560.822.033.05Figure 3: Mechanical Results for SLM 304L Standard ParametersIt can be seen here that for the sample model thicknesses from 4.0 mm down through 2.0 mm, the mechanical response is nearly the identical. These results correspond with504previous testing results from round specimens of roughly the same size and meets the tensile properties typically seen by annealed 304L. Once the model thickness is reduced to 1.0 mm or below, properties begin to signiﬁcantly reduce and a large drop in elongation to failure is present.

Once the model thickness is reduced to 1.0 mm or below, properties begin to signiﬁcantly reduce and a large drop in elongation to failure is present. From this data, it is apparent that the threshold from bulk to thin wall behavior for this parameter set occurs between a modeled wall thickness of 1.0 mm and 0.75 mm.Figure 4: 0.25 mm Sample Built with Standard Parameters Imaged at x25 (A) and x200 (B)Figure 5: 1.0 mm Sample Built with Standard Parameters Bulk Scan Pattern Imaged at x25To further investigate the cause of this transition, fractured samples were sectioned longitudinally, mounted, and imaged. The micrographs of the 0.25 mm and 1.0 mm model thickness samples can be found in Figures 4 and 5, respectively. These sectioned images expose multiple regions of reduced cross section and stress concentration, both of which can lead to failure initiation near the surface of the sample.

These sectioned images expose multiple regions of reduced cross section and stress concentration, both of which can lead to failure initiation near the surface of the sample. Even with these imperfections, the samples still exhibited a typical ductile failure as indicated by necking in the gauge section and the angled shear plane on the fractured end. Figure 5 shows grouping of internal pores near the surface of the sample. This location corresponds well to the interface between hatch ﬁll and border scans. At a thickness of 0.25 mm nearly 60% of the cross section is border scans. This phenomenon seen in the 1.0 mm sample is505likely still in eﬀect in the 0.25 mm model thickness, but any transition is hard to discern. Figure 4 shows a high amount of surface inconsistencies and voids in the 0.25 mm model thickness sample.

Figure 4 shows a high amount of surface inconsistencies and voids in the 0.25 mm model thickness sample. The gross failure in this sample is likely due to the high percentage of border parameters resulting in low interlayer strength, sub surface porosity, and and stress risers resulting from surface roughness.When the images of the fractured thin wall samples are compared to a micrograph of a 12.5 mm cube in Figure 6, it can be seen that a bulk sample in the as built condition has subsurface porosity located at the hatch ﬁll and border scan interface, but at much lower frequency.

Overall, results indicate that although these border scan parameters result in a relatively uniform surface ﬁnish, as they begin to dominate the cross section of a part, tensile properties are aﬀected.Figure 6: 304L Bulk Sample Imaged at x300Figure 7: Fractography of 0.25 mm Sample Imaged at x151 (A) and x500 (B) Showing Surface InconsistenciesFractography was performed on the samples and results for the 0.25 mm model thick- ness and the 1.0 mm model thickness samples. Figure 7 reveals several points of reduced506cross section and partially melted particles within the 0.25 mm model thickness sample. Some localized regions of ductile fractures were found as seen in Figure 8, however the fracture surface was predominantly a shear lip. For the 1.0 mm model thickness sample in Figure 9, analysis reveals a much diﬀerent surface morphology. Predominant regions of ductile fracture with a shear lip shown in Image C.

For the 1.0 mm model thickness sample in Figure 9, analysis reveals a much diﬀerent surface morphology. Predominant regions of ductile fracture with a shear lip shown in Image C. Collections of large pores around the surface as seen in the cross sectioned images are present and were measured to be approximately 20 µm in size.Figure 8: Fractography of 0.25 mm Sample Imaged at x3000 Showing Microvoid Coalescence and Shear FailureFigure 9: Fractography of 1.0 mm Sample Imaged at x150 (A) and x500 (B) Showing Subsurface Pores, and x3520 (C) Showing Shear Voids507The results for the EBM Ti-6Al-4V samples can be found in Figure 10. From these results it can be seen the the samples above 2.0 mm model thickness all preform the same. Below 2.0 mm model thickness strength begins to reduce and there is a drop in elongation at break from approximately 6% to 2%.

Below 2.0 mm model thickness strength begins to reduce and there is a drop in elongation at break from approximately 6% to 2%. The smallest three sample sets, with models at 0.25, 0.50, and 0.75 mm model thickness built at averages of 0.78, 0.79 and 0.81 mm, respectively, showing that the minimum feature size for this combination of machine and material is approximately 0.8 mm. Similar to the SLM 304L, this material exhibits a reduction of tensile properties below a critical size due to reduced eﬀectiveness of the processing parameters.200150Yield Yield UTS UTS Elongation1412) I S K (h t g n e r t S1001086E l o n g a t i o n(%)5042000.78 3.87 1.01 Average Measured Sample Thickness (mm)0.790.811.912.88Figure 10: Mechanical Results for EBM Ti-6Al-4V Standard ParametersIn order to compensate for the thin wall eﬀect seen in the SLM 304L standard scan pattern discussed previously, modiﬁcations were made.

The ratio of hatch ﬁll area to border scan area was increased by reducing the ﬁll hatch oﬀset to 0 µm and reducing the number of border passes from 2 to 1 (Figure 11). Results of the mechanical testing for the compensated parameter set can be found in Figure 12. It can be seen that although strength in not improved, the threshold of thin wall to bulk has shifted from the 1.0 mm to 0.75 mm model thickness range down to the 0.5 mm to 0.25 mm model thickness range, as indicated by the shift in the drop oﬀ of elongation at break. The 0.25 mm and 1.0 mm cross section sample images for the compensated parameter set can be found in Figures 13 and 14, respectively. Investigation of the cross sectioned samples revels an almost total elimination of the large surface porosity seen with the standard parameter set. The sample thickness is also more consistent, resulting in a reduction of stress concentration points.

The sample thickness is also more consistent, resulting in a reduction of stress concentration points. This agrees with the mechanical test results showing improved elongation to fail- ure.508Figure 11: Schematic of Compensated AM250 Border Oﬀsets100120100Yield Yield UTS UTS Elongation80) I S K (h t g n e r t S8060406040E l o n g a t i o n(%)2020000.42 4.04 1.15 Average Measured Sample Thickness (mm)0.680.922.063.06Figure 12: Mechanical Results for SLM 304L Compensated ParametersFractography analysis was performed on the samples produced with the compensated scan pattern. The 0.25 mm and 1.0 mm model thickness fracture images are found in Figures 16 and 15, respectively. Analysis of the 0.25 mm model thickness sample reveals a more uniform cross section with much larger regions of ductile failure than in the stan- dard parameter set samples. Unmelted powder is still seen, however it is limited to only the surface of the sample.

Unmelted powder is still seen, however it is limited to only the surface of the sample. Investigation of the 1.0 mm model thickness sample reveals that previously seen sub surface porosity is still present, although the pore size has been reduced to approximately 10 µm.509Figure 13: 0.25 mm Sample Built with Compensated Scan Pattern Imaged at x25Figure 14: 1.0 mm Sample Built with Compensated Scan Pattern Imaged at x25Figure 15: Fractography of 1.0 mm Sample with Compensated Scan Pattern Imaged at x300 (A) and x1000 (B) Showing Sub Surface Pores510Figure 16: Fractography of 0.25 mm Sample with Compensated Scan Pattern at x100 (A), x500 (B), and x2000 (C) Showing Ductile FailureConclusionIn this study, the eﬀect of reducing wall thickness of tensile specimens was analyzed for 304L samples built on the SLM platform and Ti-6Al-4V samples built on the EBM platform.

It was seen that by building with a parameter set based on samples representing a bulk response, mechanical properties begin to reduce once a wall thickness threshold is reached. By compensating the scan pattern to increase the area of hatch ﬁll, a more uniform sample cross section with reduced subsurface porosity has was achieved, result- ing in a shift of the threshold to a smaller thickness. However, even with a scan pattern optimized for producing thin wall geometry, a reduction in mechanical properties was still seen in parts produced with greatly diﬀerent materials by the SLM and EBM powder bed processes. This implies that the reduction in tensile properties is a generic response to the reduction of wall thickness in AM parts. Based on cross section micrographs and fractography, it appears that surface roughness and inconsistencies could be the root of this tensile property reduction.

Based on cross section micrographs and fractography, it appears that surface roughness and inconsistencies could be the root of this tensile property reduction. Further work needs to be conducted to conclude if this eﬀect is caused by the rough surface ﬁnish resulting in stress risers and surface initiated failure.511AcknowledgmentsThe Authors would like to acknowledge the assistance provided for this research by AM equipment technicians Kevin Davis and Justin Tannehill, as well as microscopists Lauren Elder and Brandon Cox.All data prepared, analyzed and presented has been developed in a speciﬁc context of work and was prepared for internal evaluation and use pursuant to that work autho- rized under the referenced contract.

Reference herein to any speciﬁc commercial product, process or service by trade name, trademark, manufacturer, or otherwise, does not nec- essarily constitute or imply its endorsement, recommendation, or favoring by the United States Government, any agency thereof or Honeywell Federal Manufacturing & Technolo- gies, LLC.This presentation has been authored by Honeywell Federal Manufacturing & Tech- nologies under Contract No. DE-NA0000622 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for pub- lication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.References[1] ASTM E8 / E8M 15a Standard Test Methods for Tension Testing of Metallic Mate-rials.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCHARACTERIZATION OF FEEDSTOCK IN THE POWDER BED FUSION PROCESS: SOURCES OF VARIATION IN PARTICLE SIZE DISTRIBUTION AND THE FACTORS THAT INFLUENCE THEMJ. Whiting, J. FoxEngineering Laboratory National Institute of Standards and Technology, Gaithersburg, MD 20899AbstractSubstantial efforts have been placed on characterizing and modeling additive manufacturing processes. The wide scope of work already done has focused on the effects of process parameters such as laser power, hatch spacing, scan speed and strategy, and layer thickness on the final part’s properties. However, the characteristics of the actual powder should also be considered.

However, the characteristics of the actual powder should also be considered. The particles’ size, morphology, roughness, and chemical composition will affect the final part properties including surface texture, density, tensile strength, and hardness. This paper will share some of the measurement methods used at the National Institute of Standards and Technology (NIST) to better understand metal powder for additive manufacturing. These include the striation/separation in transportation and handling, sampling procedures, and the actual spreading of powder in the laser powder bed fusion process.

These include the striation/separation in transportation and handling, sampling procedures, and the actual spreading of powder in the laser powder bed fusion process. Results are presented that illustrate variations in the particle size distribution as a function of location on the build platform, substrate/part surface condition, and vertical position.Keywords: Powder bed fusion; particle size distribution; powder spreading1 IntroductionThe powder bed fusion (PBF) process allows complex geometries to be additively manufactured (AM) in various metals and alloys. The process consists of a focused laser (≈ 100 µm spot size) that melts specific regions of metal powder. Between laser scans, the build area is incrementally lowered allowing the recoating mechanism (usually a stiff blade or roller) to spread thin, 20 μm to 50 µm, layers of metal powder.

Between laser scans, the build area is incrementally lowered allowing the recoating mechanism (usually a stiff blade or roller) to spread thin, 20 μm to 50 µm, layers of metal powder. While there has been increased industry use in recent years, the inability to ensure repeatable final part properties remains a primary factor preventing the widespread use of AM processes. These properties include tensile and fatigue strength, hardness, surface roughness, and density, all of which are highly dependent on the process parameters and conditions. The localized heating and cooling introduces residual stress that is often location dependent due to changes in geometry or scan strategy [1]. The environmental conditions as well as the part layout have been found to affect the final part properties [2-3].

The environmental conditions as well as the part layout have been found to affect the final part properties [2-3]. Additionally, the final mechanical properties of a part depend on the orientation in which a part is built [4].Though there has been substantial focus on documenting, modeling, and providing solutions for these known sources of variability in the PBF process, focus must also be placed on understanding the role played by the characteristics of the metal powder. Yadroitsev et al., using an analysis of variance, found the particle size distribution (PSD) to be a statistically significant1057variable affecting the contact angle of a single track and the ratio of the remelted depth to the height of a single track [5]. Other groups have conducted similar research. Spierings et al.

Other groups have conducted similar research. Spierings et al. conducted an experiment using three stainless steel powders (316L), each of different size distributions to investigate the effect of the PSD of powder on the density, surface roughness, and mechanical properties of final parts. They found particle size to have significant effect on all three of the dependent variables measured. While it was possible to create 99% dense parts using each powder, the scan strategies needed to be individually optimized for each [6]. In a similar fashion, Gu et al. investigated the microstructural and mechanical properties of AM parts created using a titanium alloy powder (Ti6Al4V) procured from three separate suppliers. Again, using equivalent parameters, powders with different PSDs produced final parts with substantial differences in their tensile properties [7]. A similar comparison is made by Liu et al.

Again, using equivalent parameters, powders with different PSDs produced final parts with substantial differences in their tensile properties [7]. A similar comparison is made by Liu et al. using two powder types with similar mean sizes (< 2 µm difference), but dissimilar distributions. One of the 316L powders had a log-normal distribution and the other featured a skewing towards smaller particles. Although at low energy densities (created by either reducing the beam size or increasing the scan speed) the log-normally distributed powder created parts with significantly lower density and substantially rougher side surfaces, the same powder also produced parts with higher tensile properties [8].All of the aforementioned findings share one thing in common: they serve as evidence that the PSD of the powder used in AM plays a significant role in determining the final part properties.

Recognizing this, it becomes apparent that any sources of variability in the PSD of AM powders must be identified and documented. Slotwinski et al. have found PSD of the AM materials to vary depending on the location in the machine [9]. They found that particles larger than 60 µm were more likely to be spread past the build plate and consequently were not used in the build. The variability was attributed to the limited space (a layer height of 20 µm was used in the experiment) between the lower edge of the recoating mechanism and the prior layer of powder. It was theorized that the larger particles were unable to easily pass under the recoater, and were therefore spread over the powder bed and eventually into the collector.

It was theorized that the larger particles were unable to easily pass under the recoater, and were therefore spread over the powder bed and eventually into the collector. Besides this work, there has been very little published that characterizes spatial inconsistencies of the PSD of AM powders.As it has been clearly shown that the PSD affects the repeatability of the PBF process and a variability in the PSD as a function of the location in the machine was documented [9], it is paramount that this spatial dependence of the PSD be investigated further. The work presented here investigates sources of variability in the PSD of metal AM powder in three situations: in a container in an as-received condition, in a loose powder bed, and in a powder bed with solidified parts.2 Experimental Procedure2.1 SamplingWhile the situation in which powder is collected may differ, the basic tools remain the same. Firstly, the powder must be extracted in a controlled manner.

Firstly, the powder must be extracted in a controlled manner. As directed in ASTM B215 [10], a slot sampler should be used to extract powder in a loose, bulk condition. The conventional slot sampler, similar to a soil core sampler, is used to collect a continuous cylindrical volume of powder, making it nearly impossible to keep a free-flowing granular material from mixing during the sampling. In order to quantify any vertical segregation of particle sizes, discrete volumes must be sampled at incrementally deeper positions. The custom slot sampler, shown in Figure 1, features a conical tip to reduce the disturbance of the bulk powder as the device is plunged into the granular material as well as both a rotating collecting shaft that allows a controlled opening or10582closing of the sampling volumes and a handle that provides ergonomic handling of the sampler.

The volumes of the collection pockets are chosen to provide an optimal sample mass (≈0.3 g) for the later-described particle size analysis. While too little powder will decrease the accuracy of the PSD by measuring less powder, alternately, if too much powder is collected, the powder must be resampled, which often introduces biases depending on the sampling location. These biases can be minimized or possibly completely avoided with the use of a proper sampling technique like a riffler. Riffling is a technique that allows a bulk powder to be divided into smaller samples, each of which is a good representation of the original powder. The technique, suggested in ASTM B215 and illustrated in Figure 2, consists of the collection of powder from a stream of powder. The bulk powder is directed towards a chute by means of vibration. Upon exiting the chute, the powder is allowed to fall into a series of moving containers.

The bulk powder is directed towards a chute by means of vibration. Upon exiting the chute, the powder is allowed to fall into a series of moving containers. If any segregation was present, it will be divided amongst each of the containers, assuming the riffler is operated correctly. With increasing numbers of rotations of the collection containers, the effectiveness of the operation increases, but the rule-of-thumb is that 100 rotations must be completed by the time all of the powder has flown through the chute.The other situation in which powder must be collected is when it is sparsely spread on a solidified AM part. The part’s upper surface is relatively rough (Ra from 9 μm to 19 µm [8]), which makes collection difficult. Mechanical removal, such as brushing or scraping, is ineffective in removing the powder. An apparatus was designed and fabricated allowing a vacuum to be used to collect powder in a controlled manner.

An apparatus was designed and fabricated allowing a vacuum to be used to collect powder in a controlled manner. This device will be described in detail in a later publication.Figure 1: Custom slot samplerFigure 2: Schematic of a spinning riffler [10]2.1.1 Powder As-receivedThree 10 kg containers of stainless steel powder (17-4PH) are used to investigate any segregation of the powder’s PSD. Most powder manufacturers ship metal powders in a plastic container that has been filled with argon prior to being sealed shut. The container often contains a silica gel desiccant to absorb any moisture. Each of the containers is opened and the desiccant10593removed prior to being sampled with the slot sampler. All containers are sampled at their radial centers at three different heights. The lowermost sample has its vertical center at a height approximately 7 mm from the bottom, while the middle and top samples are approximately 44 mm and 88.5 mm from the bottom, respectively.

The lowermost sample has its vertical center at a height approximately 7 mm from the bottom, while the middle and top samples are approximately 44 mm and 88.5 mm from the bottom, respectively. Each of the samples taken is approximately three grams, which, being too large for the particle size analyzer, is riffled into eight samples prior to being measured.2.1.2 Powder BedA loose powder bed is created by repeatedly spreading 20 µm layers of powder in a commercial PBF system without scanning the laser. The system uses a dispensing bed that moves incrementally up, providing powder to be spread across the build plate. The bed is built up to 2500 layers (50 mm total depth) prior to being sampled.

The system uses a dispensing bed that moves incrementally up, providing powder to be spread across the build plate. The bed is built up to 2500 layers (50 mm total depth) prior to being sampled. The custom slot sampler, shown in Figure 1, is used to sample powder at nine locations in the build bed (see Figure 3) and one centrally located position in the dispenser.Figure 3: Sequence and locations of powder collection; the dark outlined perimeter refers to the outer edges of the build plateFigure 4: Custom slot sampler extracting powder in position #32.1.3 Powder on a Solidified PartIn order to be able to extract powder from the surface of a solidified part, the aforementioned vacuum apparatus is used. Filters with a pore size of less than 6 µm ensure that no particles pass through the filter.

Filters with a pore size of less than 6 µm ensure that no particles pass through the filter. Powder is also extracted from before and after the part (where right refers to the side of the machine in which powder is dispensed) in order to allow the effect that the part’s presence has on the spreading of powder to be better understood. The custom slot sampler (Figure 4) is used to extract powder from the part’s vicinity. The sampler is pushed against the part’s right side so that the outer edges of each are tangential. This is done to ensure that the sample extracted was from powder that was deposited onto the loose powder bed immediately prior to the recoater passing over the part.

This is done to ensure that the sample extracted was from powder that was deposited onto the loose powder bed immediately prior to the recoater passing over the part. Powder is taken in a similar manner on the Collector10604side to ensure the sample taken is from the powder that was pushed over the part and deposited in the region to the part’s left (see Figure 5 & 6).Figure 5: Build plate and part layoutFigure 6: Sampling regions on, right, and left of solidified part2.2 Measuring Size and PSDAn optically-based particle size analyzer is used to characterize the powder. The device uses two charge-coupled device cameras to conduct a dynamic image analysis method (ISO 13322- 2). One of the cameras uses a higher magnification to resolve smaller particles and provide more accurate morphology measurements. The camera has a resolution of about 1 µm per pixel. There are numerous metrics that can be used to characterize the size of a particle, not to mention its morphology [11].

The camera has a resolution of about 1 µm per pixel. There are numerous metrics that can be used to characterize the size of a particle, not to mention its morphology [11]. When measuring a particle that is not perfectly spherical, the methods used to characterize even something as simple as the diameter are not direct. Figure 7 illustrates a few of these methods. The Martin diameter uses the area bisector of the particle’s profile, while the Feret or caliper diameter is the distance between two tangent lines that are perpendicular to the measurement direction. Another commonly used metric is the projected area diameter, which is the diameter of a circle having an equivalent area as the particle. Since nearly all of the particles in AM metal powders are nearly spherical and convex [9], the Feret or caliper diameter is reported in this paper. Additionally, the results from a size analysis can be reported in terms of the number, weight, or volume of the particles.

Additionally, the results from a size analysis can be reported in terms of the number, weight, or volume of the particles. Since the optically-based method measures the area of the particle’s projection, and this area of a nearly spherical particle can easily be converted to volume, the results will be reported in terms of volume.10615Figure 7: Methods of describing a non-spherical particle's diameter [12]Results from particle size measurements are most often reported as either a cumulative or a frequency size distribution. A frequency size distribution plot is similar to a histogram with a series of size bins on the x-axis and the percentage of their respective populations (based on number, volume, or mass) on the y-axis. A cumulative size distribution also has the size of the particle on the x-axis, but the y-axis is displayed in terms of the accumulative percentage of the particles.

A cumulative size distribution also has the size of the particle on the x-axis, but the y-axis is displayed in terms of the accumulative percentage of the particles. As mentioned, the metric of “particle size” can be specified in many ways; this paper will report the particle size in terms of the Feret diameter. The cumulative and frequency distribution will always be reported in terms of particle volume. The D10, D50, and D90 of a sample’s PSD refer to the maximum size corresponding to 10 %, 50 %, or 90 % of the particles. Unless otherwise noted, each powder sample is measured five times, and the plotted PSDs use the average of these five measurements for each size class with the standard deviation plotted as error bars.3 Results and Discussion3.1 Powder As-receivedTable 1 and Figures 8-10 illustrate the PSDs at the three heights of the three containers sampled. It should be noted that the containers sampled were known to have dissimilar PSDs.

It should be noted that the containers sampled were known to have dissimilar PSDs. To show this, the frequency distribution of the top location of each powder is plotted in Figure 11. Powder B has virtually no segregation of particle sizes, while Powder A and Powder C seem to have a slight variability with height. The sample taken from the middle height has the smallest particles of all three of the powders, though as shown in Figure 9, the change in the PSD of Powder B’s samples is almost indiscernible. From this result, the difference in the initial PSDs seems to affect the segregation seen in the containers.

The blue outline represents the 25th and 75th percentiles of the five measurements taken while the red line shows the median. The ‘whiskers’ (referring to the black lines) are the range of all values that are not considered outliers. A datum is considered an outlier if it is larger than q3 + 1.5(q3 – q1) or smaller than q1 – 1.5(q3 – q1), where q1 represents the 25th quartile and q3 the 75th. Figure 12, as well as Table 1, show the difference between the three locations. The powder sampled from the ‘middle’ location is smaller in all three powder types, but the difference is much more prevalent in Powder C and Powder A. A similar trend is seen in the D50 and D90 values. This vertically centered concentration of smaller10637particles may be due to the manner in which the powder was loaded into the container or the disturbance from handling and transportation.

This vertically centered concentration of smaller10637particles may be due to the manner in which the powder was loaded into the container or the disturbance from handling and transportation. Jenike theorized [13], Rémi and Meakin modeled [14], and Savage and Lun showed empirically [15] that finer particles often concentrate near the center of a pile of powder that has been poured. The smaller particles tend to cohere to one another, while the larger particles roll off to the side more easily. It is possible this same mechanism is driving the segregation seen here. As the container is rotated and jostled, the larger particles may move more easily to the extremities of the container. These small variations in the PSD as a function of height likely do not warrant the need for mixing prior to use.

These small variations in the PSD as a function of height likely do not warrant the need for mixing prior to use. It is noted that more research is needed to verify these findings, as this is a narrow view of this segregation.Figure 12: Box plots including 25th and 75th percentiles, median, range and outliers of D10 values of maximum Feret diameter; values are volume-based and data points are considered outliers if larger than q3 + 1.5(q3 – q1) or smaller than q1 –1.5(q3 – q1)3.2 Powder Bed SamplingShown in Figure 13, there is very little difference in the powder sampled across the loose powder bed. The D50 of the powder sampled shows the largest range of size at 1.5 µm, which is very close to the 1 µm resolution of the optical particle size analyzer. As the variability of the PSDs approach this resolution, concerns regarding the uncertainty introduced throughout the experiment arise.

As the variability of the PSDs approach this resolution, concerns regarding the uncertainty introduced throughout the experiment arise. Furthermore, the standard deviation from the five measurements used in calculating the PSDs ranges from 0.09 µm to 0.74 µm (Table 2) which is of a similar magnitude as the ranges documented. These trends are representative of what is seen in other samples.10648Figure 13: Cumulative size distribution (volume-based maximum Feret diameter) of powder sampled from a loosely spread powder bed using the custom slot samplerTable 2 lists the standard deviations (σ) and D10, D50, and D90 of each of the powder samples from the ‘7’- ‘8’- ‘9’ row (see Figure 3 for collection locations). The trends seen here are indicative of a spreading process that has little influence on the trends previously seen in the PSD of powder in the PBF process. This is potentially due to the loose, unpacked nature of the powder bed.

This is potentially due to the loose, unpacked nature of the powder bed. While a limited gap between the recoating blade and the part surface or substrate limits the size of particles that can pass under, a loose powder bed may allow larger particles to embed themselves into the previously spread particles.

The next section explores the influence of solidified parts on the PSD of the spread powder.Table 2: D10, D50, D90 and standard deviations of the maximum Feret diameters and respective locations (diagram in Figure 3)D10D50D90Powder Location7 8 9 Dispenser (center)Average (μm) 22.40 21.84 22.78 22.72σ (μm)0.12 0.09 0.19 0.08Average (μm) 34.86 34.38 35.88 35.78σ (μm)0.17 0.22 0.36 0.16Average (μm) 53.56 54.54 55.00 54.661065σ (μm)0.29 0.51 0.19 0.7493.3 On and Off Part SamplingAs seen in the previous two sections, there is minimal segregation of particle sizes in the as-received powder and in a spread powder bed, but there has been documented variability in the PSD of powder in the collector versus powder in the dispenser [9]. Therefore, the presence of solidified part surfaces must play a significant role in the segregation of various particle sizes.

Therefore, the presence of solidified part surfaces must play a significant role in the segregation of various particle sizes. Figure 14 contains a cumulative size distribution of the powder sampled from the surface as well as from the right and left of a solidified part. As previously mentioned, the right side refers to the side of the part closer to the dispenser. An obvious skew of the size distribution is seen between the powder taken from the part and samples taken from its vicinity. The ‘Part’ powder is substantially smaller, while the ‘Right’ powder is only slightly (a range of D50 of 1.16 µm) smaller than the ‘Left’. There is also a change in the distribution of the ‘Part’ powder, which is more easily visualized in a frequency distribution (shown in Figure 15). The distributions of the ‘Left’ and Right samples are similar with a slight skew towards smaller particles on the Left.

The distributions of the ‘Left’ and Right samples are similar with a slight skew towards smaller particles on the Left. The PSD of the Part powder has substantially smaller particles (D50 =31.54 µm compared to 35.08 μm and 36.24 µm for the ‘Left’ and ‘Right’ samples, respectively) and is narrower. The surface of the solidified part appears to act as a filter of the powder being spread. While the step in the layer height is only 20 µm, due to the densification of the loose powder during the melting and solidification, the actual vertical distance between the recoater blade’s underside and the solidified part surface is substantially larger. Using a similar method as theorized in [16], the effective, steady-state layer height for a 20 µm build plate movement is approximately 33 µm, assuming a 99% dense final part. Since this is around the D50 (maximum Feret diameter) of the powder spread, a large portion of particles are unable to pass under the recoating blade.

Since this is around the D50 (maximum Feret diameter) of the powder spread, a large portion of particles are unable to pass under the recoating blade. Though due to the relatively rough top surface of the solidified part, the actual effective layer thickness is highly variable and might allow larger particles to settle in the valleys between laser tracks.Figure 14: Cumulative size distribution (volume-based maximum Feret diameter) of powder sampled from the left, right, and surface of a solidified part106610Figure 15: Frequency size distribution of powder sampled from the left, right, and surface of a solidified part; Red = on part, Blue = right of part, Green = left of part4 ConclusionsThe powder spreading in the PBF process is extremely complex. The fact that each powder particle is unique and that the cohesion forces between particles is dependent on the surfaces and geometries of these particles creates a very intricate process.

The fact that each powder particle is unique and that the cohesion forces between particles is dependent on the surfaces and geometries of these particles creates a very intricate process. Moreover, the intricate surfaces of solidified parts have been found to influence the spreading. While simplifications can and must be made in modeling efforts (e.g., perfectly spherical particles, smooth surfaces, finite size ranges), empirical analysis of the powder spreading must be conducted to ensure that the correct assumptions are made. The work presented here delineated the segregation of certain particle sizes in three situations. A minimal change in the PSD at different heights of the powder in its as- received condition was seen. There was also very little change in the size of the powder at different locations along the recoating direction of a loose powder bed.

There was also very little change in the size of the powder at different locations along the recoating direction of a loose powder bed. Significant differences in the particles collected from the top of a part versus the part’s adjacent vicinity was seen, and potential reasons for this have been proposed. It is noted that the work shown here provides only a glimpse into the spatial dependence of the PSD of powder and should not be taken as concrete evidence but rather as a starting point for future research. Further research will focus on closely investigating the interaction between the powder and the recoating mechanism via empirical analysis as well as modeling. The rheological behavior of AM powders will be compared to the performance of the spreading in the PBF process.106711References[1] Mercelis, Peter, and Jean-Pierre Kruth. "Residual stresses in selective laser sintering andselective laser melting." Rapid Prototyping Journal 12, no.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCharacterization of Heat-Affected Powder Generated during the Selective Laser Melting of 304L Stainless Steel PowderAustin T. Sutton1, Caitlin S. Kriewall2, Ming C. Leu1, Joseph W. Newkirk21Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409 2Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractThe selective laser melting (SLM) process is an Additive Manufacturing (AM) technique that uses a laser to fuse successive layers of powder into near fully dense components. Due to the large energy input from the laser during processing, vaporization and instabilities in the melt pool occur causing the formation of condensate and laser spatter, collectively known as heat-affected powder.

Since heat-affected powder settles into the powder bed, the properties of the unconsolidated powder may be altered compromising its reusability. In this study, characterization of 304L heat- affected powder was performed through particle size distribution measurements, x-ray diffraction, metallography, energy-dispersive spectroscopy mapping, and visualization of grain structure with the aid of a focused-ion beam. The results show morphological, microstructural, and surface chemistry differences between the starting powder and heat-affected powder formed during processing which aid in the understanding of laser spatter and condensate that form in the SLM process.IntroductionPowder-bed fusion refers to a subset of Additive Manufacturing (AM) processes that bond powder layers in a sequential manner for the production of three-dimensional components directly from Computer-Aided Design (CAD) files [1].

Among this class of manufacturing techniques is the selective laser melting (SLM) process where a laser scans across a powder-bed melting and consolidating particles to form parts that approach theoretical density [2,3]. The layer thickness is often no more than 70 µm facilitating the construction of fine features and, thus, overall geometrical complexity. Consequently, SLM has attracted a significant amount of interest from a wide variety of areas involving biomedical, aeronautical, and automotive applications.A large portion of the research conducted in SLM has been focused on adjusting laser process parameters for manufacturability of a variety of materials such as Ti-6Al-4V [4,5], 316/316L stainless steel [6,7], and IN718 [8,9]. Typically, these variables include the laser power, scan speed, and hatch spacing for which an optimal combination is found to minimize porosity.

Typically, these variables include the laser power, scan speed, and hatch spacing for which an optimal combination is found to minimize porosity. Other work in SLM has been centered on the interaction of the laser beam with the powder-bed through heat and fluid flow simulations for insight into melt pool formation and cooling rates [10– 12]. In general, it is found that the melt pool experiences high thermal gradients giving rise to strong Marangoni convection currents, which play a large role in the wetting behavior of the melt on the previously solidified layer. Adding to the complexity of the process is the dependence of the melt pool wettability on the oxygen content where the balling phenomenon can be induced as261a result of the presence of oxides [13,14]. The onset of balling indicates a lack of wetting and can ultimately cause the formation of porosity and possible damage to the recoating mechanism.

The onset of balling indicates a lack of wetting and can ultimately cause the formation of porosity and possible damage to the recoating mechanism. Results from simulations also show that the melt pool experiences a high degree of superheat leading to vaporization of material and subsequent ejection of material. Rapid vaporization of material is known to cause melt pool instabilities by providing a large upsurge of vertical momentum in the melt pool causing molten material to be ejected in the form of laser spatter [15– 17]. Moreover, vaporized material that escapes the melt pool rapidly solidifies in the chamber atmosphere causing the formation of nano-particles, more commonly known as condensate, that coat the inside of the build chamber during the SLM process.

If generated in excess or not adequately removed, condensate can lead to attenuation and defocusing of the laser beam causing pore formation due to a lack of fusion [18–21].The collection of both laser spatter and condensate is hereby referred to as heat-affected powder as it is formed under high temperature conditions originating from the melt pool. In SLM, studies regarding laser spatter characterization and the potential consequences of its formation on part quality are seldom. Simonelli et al. [22] characterized the spatter formed from 316L stainless steel, Al-Si10-Mg, and Ti-6Al-4V generated during SLM. It was found that laser spatter is not only comprised of particles much larger than the base powder, but is also chemically different as a result of oxidation on the outer surfaces of the particles. Liu et al. [19] also found that laser spatter generated from processing 316L was large with some particles having an approximate diameter of 400 µm.

Liu et al. [19] also found that laser spatter generated from processing 316L was large with some particles having an approximate diameter of 400 µm. Although typical procedures involve sieving of used powder to eliminate these large particles for recycling, Liu et al. processed used 316L powder without sieving and found that the tensile properties of the parts exhibited decreased strength and elongation which were attributed to laser spatter contamination of the powder-bed. From both of these studies, it is clear that laser spatter’s large size allows for it to settle into the powder-bed thus compromising the reusability of the powder.

From both of these studies, it is clear that laser spatter’s large size allows for it to settle into the powder-bed thus compromising the reusability of the powder. These large particles as well as the formation of oxides on their surfaces may be a reason why researchers often find a coarsening particle size distribution and increase in oxygen content when performing recycling studies for various materials [23–25].Although characterization of laser spatter has been performed in the aforementioned studies, more in-depth research needs to be conducted to rigorously understand the properties of laser spatter and its potential impact on the SLM process. Since laser spatter and condensate are coupled, characterization of condensate is also necessary, which, to our knowledge, has yet to be considered in the SLM community. Therefore, this study characterizes heat-affected powder in an attempt to understand how 304L stainless steel evolves with reuse during the SLM process.

Therefore, this study characterizes heat-affected powder in an attempt to understand how 304L stainless steel evolves with reuse during the SLM process. The characterization performed highlights the morphological, chemical, and microstructural differences observed between virgin 304L and heat-affected powder, thereby providing explanations for the changes often observed by researchers in recycled powderExperimental MethodsA Renishaw AM250 SLM machine was used for fabrication of parts whose sole purpose was to produce heat-affected powder for subsequent characterization. The AM250 is equipped with a 200W Nd-YAG 1070 nm pulsed laser that has a Gaussian beam intensity profile. To ensure minimal oxidation of the parts, the entire build chamber was rendered inert until a stabilized oxygen content below 1000 ppm was observed.

To ensure minimal oxidation of the parts, the entire build chamber was rendered inert until a stabilized oxygen content below 1000 ppm was observed. During operation, a constant 400 ft3/min262volumetric flow rate of argon crossflow was maintained across the build plate to serve as a shielding gas against oxidation and for adequate removal of melt pool ejecta from the path of the laser beam.Since heat-affected powder is formed by the interaction of the laser beam with the powder- bed, increasing the area to be melted in each layer ultimately leads to a larger amount of spatter and condensate generation. Thus, in order to generate a significant amount of heat-affected powder for analysis, it was desired to utilize a large portion of the build area without requiring copious time and resources.

Thus, in order to generate a significant amount of heat-affected powder for analysis, it was desired to utilize a large portion of the build area without requiring copious time and resources. Consequently, 50% of the substrate area was chosen as a means to fulfill these requirements where a 5x5 array of plates each 15 mm in height was selectively melted as illustrated in Figure 1.The powder used in this study was argon gas-atomized 304L stainless steel purchased from LPW Technology with its chemical composition listed in Table 1. Before being inserted into the Renishaw AM250, all virgin powder was sieved using a 63 µm screen in an argon atmosphere for removal of aggregates, breakup of agglomerates, and minimization of oxygen introduction into the build chamber.

Once loaded in the SLM machine, the powder was dispensed and spread across the substrate for assessment of layer uniformity.Due to direction of the argon crossflow, the location immediately downstream of the build plate (Figure 1) becomes concentrated in laser spatter and condensate during part fabrication. For this reason, a powder sample was taken from this area (heat-affected powder) using a clean spatula, thereby excluding the unconsolidated powder in the powder bed. An additional sample was taken from one of the walls of the build chamber in attempt to obtain a highly concentrated sample of condensate material. Characterization of this sample was performed with the use of a scanning electron microscope (SEM) as well as energy-dispersive spectroscopy (EDS).

Characterization of this sample was performed with the use of a scanning electron microscope (SEM) as well as energy-dispersive spectroscopy (EDS). In an attempt to separate laser spatter from condensate for characterization of each individually, the heat-affectedIndividual PartHeat-Affected Powder Collection Site30 mm30 mm15 mmSubstrateInert Gas Flow DirectionFigure 1. Illustration of the experimental setup of the build. The combined area of the parts for each layer corresponds to 50% of the substrate area.263powder was sieved using an ATM Sonic Sifter using 25 µm and 45 µm screens for separation into three size ranges, namely, -25 µm, +25 -45 µm, and +45 µm. Each size fraction was then homogenized using a Turbula T2C mixer to eliminate sampling bias and analyzed in regards to morphology, chemistry, and microstructure.Table 1.

Each size fraction was then homogenized using a Turbula T2C mixer to eliminate sampling bias and analyzed in regards to morphology, chemistry, and microstructure.Table 1. Chemical composition of virgin 304L stainless steel powder.Wt %CCr0.018 18.4Cu< 0.1Element NiFe Mn NOPBal1.4 0.06 9.8 0.02 0.012SSi 0.005 0.63Morphological characterization carried out in this study employed the use of three SEMs: Hitachi S4700, FEI Helios NanoLab 600, and an ASPEX 1020. While imaging and EDS were performed across all instruments, the automated feature analysis (AFA) capability of the ASPEX 1020 was employed for generation of particle size distributions and determination of particle shape using the circularity shape factor as outlined by Cox [26]. AFA allows the user to establish a region of interest on a sample and then measure specific properties of the encompassed features including projected area, perimeter, and chemistry.

AFA allows the user to establish a region of interest on a sample and then measure specific properties of the encompassed features including projected area, perimeter, and chemistry. Each of the particle size distributions obtained was from at least 2500 particles to generate reliable distributions [27]. All number distributions were converted to a volume basis assuming a spherical shape for each particle.Chemical characterization of each powder sample involved the use of EDS for an approximation of elemental concentration. To avoid possible discrepancies in the EDS results due to particle curvature [25], each powder sample was mounted in epoxy, ground, and polished to 0.05 µm using a small force in each step to minimize particle pullout. Before insertion into the SEM, an Au/Pd coating was applied to each mounted powder sample and electrically grounded with copper tape to eliminate charging and beam deflection.

Before insertion into the SEM, an Au/Pd coating was applied to each mounted powder sample and electrically grounded with copper tape to eliminate charging and beam deflection. EDS measurements were taken at various locations on several particles for comparison of chemistry. Further use of EDS involved mapping across the polished surface of a large heat-affected powder particle for insight into potential segregation of Cr and Ni.

Further use of EDS involved mapping across the polished surface of a large heat-affected powder particle for insight into potential segregation of Cr and Ni. Oxides on heat-affected powder were investigated by a focused- ion beam (FIB) for milling and imaging the oxide/metal interface.For identification of phases present, a Panalytical X’Pert Pro Multi-Purpose Diffractometer was utilized to perform subsequent analysis using Rietveld refinement for quantification of the volume percentage of phases present.Results and DiscussionMorphological CharacterizationElectron micrographs of the -63 µm virgin 304L used for part fabrication as well as the heat-affected powder downstream of the build area sieved at various size fractions are shown in Figure 2. Although large particles greater than 100 µm clearly exist, it is apparent that particles below 25 µm exist as well.

Although large particles greater than 100 µm clearly exist, it is apparent that particles below 25 µm exist as well. Thus, heat-affected powder is comprised of particles with a large size range, all of which have the potential to be redeposited on parts or in the powder bed. Particles larger than the base powder that fall onto parts can cause issues during fabrication including lack264of fusion due to the inability to be fully melted by the laser, and damage to the recoater mechanism affecting the powder-bed density. If redeposited in the powder bed and not subsequently melted, the size distribution of the unconsolidated powder will undoubtedly be altered. As a result, the deposition of large particles in the powder bed can contribute to the coarsening of powder with continual reuse.

As a result, the deposition of large particles in the powder bed can contribute to the coarsening of powder with continual reuse. Although sieving may serve as an effective means towards removing these large particles, such a procedure will not be able to eliminate all heat-affected powder since it is composed of particles similar in size to virgin 304L.Although aggregates are present in all size fractions, it is important to note that many of the particles exhibit a high degree of circularity when compared to the base powder. Thus, the cooling rate was low enough to allow surface tension forces to shape the molten material into spheres during flight. This is the reason why the particles in +45 µm size fraction appear to be the most circular of all the samples observed. Verification of this observation was made through the quantification of the particle shape through circularity.

Verification of this observation was made through the quantification of the particle shape through circularity. The average circularity for each sample is100 µm100 µm(a)(b)100 µm100 µm(c)(d)Figure 2. Heat-affected powder sieved at various size fractions taken with ASPEX SEM, including (a) Virgin 304L, (b) -25 µm, (c) +25/-45 µm, and (d) +45 µm.265tabulated in Table 2. The closer the circularity value is to unity, the more circular the geometry. The coarsest size fraction can be seen to have the highest circularity illustrating its sphericity in comparison to the other samples. Surprisingly, virgin 304L exhibits the lowest circularity even though no aggregates or agglomerates in the sample were noticed.Table 2. A listing of the D10, D50, and D90 on a number and volume basis for each measured size distribution with incorporation of the minimum and maximum particle size.

A listing of the D10, D50, and D90 on a number and volume basis for each measured size distribution with incorporation of the minimum and maximum particle size. A comparison of shape using the circularity shape factor is also provided.SampleVirgin 304L25 µm+25 -45 µm+45 µmDn10 (µm)16.116.931.446.1Dn50 (µm)21.422.137.955.7Dn90 (µm)31.429.046.178.2Dv10 (µm)19.219.033.449.5Dv50 (µm)27.525.540.866.7Dv90 (µm) Minimum (µm) Maximum (µm) Circularity38.3 0.3 54.5 0.8133.7 1.8 48.4 0.8951 10.2 65.9 0.8394.6 20.3 123.1 0.90Figure 3 gives the particle size distributions of each sample in number and volume representations. Information regarding the D10, D50, and D90 for the number and volume distributions is given in Table 2. Among the distributions shown, the virgin 304L has the highest number of fines present even when compared to the -25 µm size fraction.

Among the distributions shown, the virgin 304L has the highest number of fines present even when compared to the -25 µm size fraction. Since one of the main goals of sieving was to separate laser spatter from condensate, this result indicates that the presence of condensate in the region where the heat-affected powder was collected is either minimal or unable to be measured by AFA due to its small size of approximately 100 nm for mild steel. However, it is important to note that particles of this size were never observed in any of the micrographs taken of the heat-affected powder samples, with the smallest particle size measured being 1.8 µm. This could indicate that a significant amount of condensate generated is entrained in the argon flow across the build area and subsequently removed rather than depositing in the powder bed.

This could indicate that a significant amount of condensate generated is entrained in the argon flow across the build area and subsequently removed rather than depositing in the powder bed. A word of caution concerning this result is that the removal of condensate strongly depends on the inlet gas flow geometry and the magnitude of the flow during the process [18,20]. Therefore, this finding does not apply universally and should be investigated by each researcher to ensure adequate removal of laser byproduct for quality control.266(a)(b)Figure 3.

Therefore, this finding does not apply universally and should be investigated by each researcher to ensure adequate removal of laser byproduct for quality control.266(a)(b)Figure 3. Particle size distributions of virgin 304L and size fractions of heat-affected powder in (a) number distribution and (b) volume distribution.While condensate was not detected in the heat-affected powder collected downstream of the build area, images of the sample taken from the chamber wall can be seen in Figure 4, which reveals the appearance of powder similar in size to what is reported in literature regarding condensate as previously mentioned. As a consequence of its size, the powder agglomerates and loosely adheres to surfaces in the build chamber producing a thin film of material that collects on most surfaces in the build chamber during processing, and exhibits a blackened appearance.

The relatively small size of condensate particles indicates that vaporization of material occurred while parts were being built and was immediately quenched in the chamber atmosphere.10 µm1 µmFigure 4. Images of a sample taken from the chamber wall after part fabrication showing heavy concentration of condensate.Chemical CharacterizationWhile Figure 2 illustrates the morphological differences between the base and heat-affected powder, it is obvious that the surfaces of some heat-affected particles contain dark regions2676582371425 µm25 µm(a)(b)Figure 5. Electron micrographs of cross-sectioned (a) virgin 304L and (b) heat-affected powder. The white arrows indicate surface features present on heat-affected particles, and the black arrows represent the location of EDS performed on said regions.indicative of a lower atomic number.

The white arrows indicate surface features present on heat-affected particles, and the black arrows represent the location of EDS performed on said regions.indicative of a lower atomic number. Since the precursor powder does not contain such features, it was of interest to chemically identify these regions and provide rationale for their formation.Identification of the chemical species present was performed by EDS in the ASPEX 1020. Rather than simply do a point EDS on the surface of particles, samples of virgin 304L and heat- affected powder were mounted in epoxy and cross-sectioned before insertion into the SEM. Figure 5 shows the powder cross-sections produced by mounting and polishing as well as the points of exposure for EDS measurement. Immediately noticeable are the previously observed structures (white arrows) present on the particle surfaces in Figure 5b where no distinct penetration into the underlying particle bulk is obvious.

Immediately noticeable are the previously observed structures (white arrows) present on the particle surfaces in Figure 5b where no distinct penetration into the underlying particle bulk is obvious. EDS measurements for each of the indicated locations are provided in Table 3. Given the uncertainty in EDS, the slight variability in the elemental concentrations in points 1 – 6 show no significant differences in bulk chemistry among the two powders. However, points 7 and 8 reveal a strong presence of Mn, Si, and O. Since Mn and Si are strong oxide formers and the observed features are located on the surface of the particles, these regions are most likely oxides that formed while the particles were in flight after being ejected from the melt pool. Simonelli [22] found a very similar result when studying 316L laser spatter formed during SLM. Chasoglou [28] also observed the same oxide island features on a water- atomized steel powder for use in powder metallurgy.

Chasoglou [28] also observed the same oxide island features on a water- atomized steel powder for use in powder metallurgy. After using XPS, it was concluded that these oxides contain Fe and Cr in addition to Mn and Si. For a more thorough characterization of the oxides present, a similar approach will need to be taken to enable a look into the bonding states of the elements present.268Table 3. EDS measurements at various locations on virgin powder and heat-affected particles.LocationFeCrElement (Wt %) MnNiOCSi1 2 3 4 5 6 7 864.7 63.0 63.4 65.3 64.2 63.0 2.6 1.522.3 22.0 21.4 21.1 21.6 21.6 7.7 5.97.0 7.9 7.5 7.5 6.7 8.4 0.5 0.52.5 2.1 3.2 2.7 3.0 3.0 17.3 9.01.9 3.3 2.3 1.6 2.5 2.2 53.4 61.61.3 1.5 1.7 1.6 1.5 1.4 7.1 5.80.3 0.3 0.5 0.2 0.5 0.4 11.3 15.825 µm3 µmFigure 6.

(a) Electron micrograph of a polished heat-affected particle with a large oxide on its surface where the black rectangle indicates a region for subsequent FIB analysis, and (b) FIB image of the oxide/metal interface, as denoted by the black arrow. While the presence of oxides on the surfaces of heat-affected particles is reasonable due to the 1000 ppm oxygen content of the build atmosphere during operation, the formation of islands of oxides is less intuitive and deserves an explanation. The current theory regarding oxidation of metals can be categorized into four stages: induction, nucleation, growth, and coalescence [29,30]. Induction is the time preceding nucleation of oxide islands where the dissociation of oxygen molecules and subsequent chemisorption of oxygen occur on the metal surface. Once a sufficient oxygen concentration is reached, stable oxide nuclei form and grow.

Once a sufficient oxygen concentration is reached, stable oxide nuclei form and grow. During the growth stage, the oxides will merge and coalesce becoming larger to eventually create a continuous film of oxide.The oxide islands in Figure 5b seem to follow this four stage progression where the features were in the growth phase until the particle cooled. One important detail of this reaction sequence is the penetration of oxide into the underlying metal as a result of oxygen diffusion into the metal lattice. For observation of consumption of the metal particle by the oxide, a FIB was used to269generate a cross-section of the oxide/metal interface of a polished heat-affected particle. Figure 6a shows the heat-affected particle selected as well as the region of interest denoted by a black rectangle.

Figure 6a shows the heat-affected particle selected as well as the region of interest denoted by a black rectangle. The oxide/metal interface in Figure 6b clearly shows a curvature meaning the oxide consumed a portion of the metal, a result that is explained in detail by Zhou [30].Due to the relatively slow cooling rate of large particles, the developed microstructure could possibly chemically segregate as a consequence of solute rejection from the melt. Depending on the solidification mode of the stainless steel, the chromium and nickel will either be retained in the solid or rejected ahead of the solidification front. Although such a result is uncommon for high cooling rates, solute redistribution is a common occurrence in alloys as the cooling rate decreases. In order to evaluate if microsegregation of elements occurs in the larger particles, EDS maps of cross-sectioned particle surface were generated.

In order to evaluate if microsegregation of elements occurs in the larger particles, EDS maps of cross-sectioned particle surface were generated. The EDS maps generated for a large heat-affected particle can be seen in Figure 7. Based on the results, microsegregation could not be detected given the resolution limit of EDS.The chemistry of the condensate was also investigated with EDS for insight into possible elemental vaporization. For this investigation, the powder sample collected from the chamber wall was examined since it was found that the powder collected there has a concentration of condensate relative to the heat-affected powder collected downstream of the build area. In order to avoid the effects of interaction volume skewing the measurements, a particle with a large amount of agglomerated condensate was found and analyzed, as depicted in Figure 8. The EDS results in Table 4 reveal a significant amount of each element found in the base powder with the exception of Ni.

The EDS results in Table 4 reveal a significant amount of each element found in the base powder with the exception of Ni. This indicates vaporization of all the elements present within the melt pool suggesting that a large amount of superheat is generated during processing. Although it is still not clear whether condensate is deposited inside the build area since none was found in the heat-affected powder sample, this result does show vaporization of all elements occurred. Therefore, potential chemistry modifications may need to be made to the precursor powder to have more control over part chemistry.FeCrCSi10 µmNi25 µmFigure 7. EDS maps of a large heat-affected powder particle showing no evidence of microsegregation.270OMn5 µmFigure 8. Heat-affected particle with condensate on its surface. The white cross-hairs indicate the location of EDS measurement. Table 4.

Heat-affected particle with condensate on its surface. The white cross-hairs indicate the location of EDS measurement. Table 4. EDS results of condensate showing a large amount of all elements present.ElementCOSiCrMnFeNiWt %12.52% 17.37% 16.53% 12.53% 5.99% 32.03% 3.02%Microstructural CharacterizationIdentification of the phases present within the powder samples was performed through XRD for virgin 304L, heat-affected powder and its size fractions. Figure 9 shows the obtained diffraction patterns indicating a strong presence of austenite and delta ferrite in all analyzed samples. Virgin 304L is nearly pure austenite while heat-affected powder shows significant delta ferrite. Moreover, the amount of each phase varies as a function of the size fraction where the primary austenite peak at 44° is less prominent as the particle size increases indicating an increase in delta ferrite.

Quantification of the XRD diffraction patterns was performed through Rietveld refinement and can be seen graphically in Figure 10 accompanied by the tabulated results in Table 5. Since the chemistry of each fraction is nearly the same as previously confirmed by EDS, the observed trend is indicative of a cooling rate difference between the smallest and largest particles. Therefore, further investigation must include the estimation of cooling rate as a function of particle size to view this correlation.271δ γγδγδ200001500025 µm) . u . a ( y t i s n e t n I100005000+25/-45 µm+45 µmVirgin 304LHeat-Affected030405060 2θ708090Figure 9. XRD diffraction patterns of heat-affected powder and its size fractions compared to virgin 304L.)%(e g a t n e c r e P e s a h P100 90 80 70 60 50 40 30 20 10 0Virgin 304L Heat-Affected25 µm+25 -45 µm+45 µmAusteniteDelta FerriteFigure 10.

Quantification of XRD diffraction patterns using Rietveld refinement illustrating the dependence of delta ferrite formation on the particle size.Table 5. Quantification of austenite and delta ferrite fractions for heat-affected powder and its various size fractions in comparison to virgin 304L.SampleAustenite (%) Delta Ferrite (%)Virgin 304L 98.5 1.5Heat-Affected 60.8 39.225 µm 93.9 6.1+25 -45 µm 88.3 11.7+45 µm 71.8 28.2272The difference in the relative proportion of austenite and delta ferrite formation can be explained through the determination of solidification mode, or the sequence of phase transformations that occurs during solidification and subsequent cooling as a solid. Since chromium and nickel are strong delta ferrite and austenite stabilizers, respectively, the solidification mode in stainless steels is often determined through the ratio of chromium to nickel equivalencies (Creq/Nieq).

A listing of the available methods for evaluation of Creq/Nieq is given by Kiorinko et al. [31]. In general, there exist four types of solidification modes [32]:1. A mode: L  L+ γ  γ for Creq/Nieq < 1.25 + γ  γ + 2. AF mode: L  L+ γ  L+  L+ 3. FA mode: L  L+ + γ  γ + 𝛿𝛿 𝛿𝛿 + γ  γ for Creq/Nieq > 1.95  4. F mode: L  L+ 𝛿𝛿 𝛿𝛿 𝛿𝛿 Using the chemistry of the virgin powder provided by LPW Technology (Table 1) and calculating the Creq/Nieq using the Hull expression, we find that 304L will solidify as primary ferrite and then austenite (FA). At high cooling rates, the solid-state transformation of delta ferrite to austenite is hindered thus allowing retention of the delta ferrite phase. However, at even higher cooling rates approaching 105 K/s, it is possible for the solidification mode to change from FA to AF as a result of a sufficient amount of undercooling.

However, at even higher cooling rates approaching 105 K/s, it is possible for the solidification mode to change from FA to AF as a result of a sufficient amount of undercooling. Since virgin 304L is gas-atomized where typical cooling rates are on the order of 105 K/s, it is reasonable that 304L is primarily austenite. If the percentage of delta ferrite decreases with increasing cooling rate, it is reasonable to assume that the smaller size fractions (-45 µm) are approaching an AF solidification mode whereas the large particles are solidifying as FA. A few studies [33,34] have demonstrated that powder properties can influence phase formation in as-built components. While the influence of particle microstructure on part properties is out of the scope of this paper, it is important to understand the formation of delta ferrite indicates the presence of heat-affected powder.

Thus, phase quantification could prove as a viable approach towards assessing the degree of which a powder has been used. γ for 1.25 < Creq/Nieq < 1.48  γ for 1.48 < Creq/Nieq < 1.95 𝛿𝛿𝛿𝛿𝛿𝛿ConclusionsIn this study, the morphological, chemical, and microstructural properties of heat-affected powder were investigated for insight into the expected changes that are to occur as a result of recycling 304L stainless steel in the SLM process. While previous studies primarily focus on laser spatter, emphasis in the current study was placed on the combination of laser spatter and condensate for a more comprehensive characterization of ejecta originating from the melt pool. Clear identification of both laser spatter and condensate was accomplished by collecting two samples: heat-affected powder downstream of the build area, and a sample from the wall of the chamber after building parts.

The collected heat-affected powder was found to be primarily laser spatter, suggesting that most of the condensate does not deposit in the build area. However, such a conclusion is highly dependent on the magnitude of argon crossflow employed as well as the geometry of the build chamber and inlet gas manifold. Thus, each researcher must study the flow inside the build chamber for insight into condensate deposition inside the build chamber. The size of laser spatter particles was found to exist across a large size range where, once ejected, deposition into the build area is possible, potentially compromising powder reuse. Although the size range is273large, the average circularity of the laser spatter in each size fraction was found to be greater than that of that the base powder. It is important to note that aggregates were found to exist as a consequence of particles merging while molten.Investigation of powder chemistry led to the observation of oxide islands on laser spatter.

These oxides are rich in Mn and Si, which are strong oxide formers. Further analysis revealed that these oxides nucleate and grow on the surfaces of particles. Depending on the growth rate of the oxides, the coalescence of these islands will vary, producing oxides of various size. EDS also allowed for identification of the chemical species available in condensate powder. Based on these results, it is clear that the condensate is comprised of appreciable amounts of all alloying elements present in the base powder, providing evidence for large superheat in the melt pool.Microstructural analysis of the heat-affected size fractions and virgin 304L through XRD showed varying amounts of austenite and delta ferrite. In general, it was found that the percentage of delta ferrite increased with increasing particle size.

In general, it was found that the percentage of delta ferrite increased with increasing particle size. These results were attributed to the large difference in cooling rates experienced by the large and small particles, enabling the switch from FA to AF solidification mode through an increased undercooling. Since delta ferrite is a marker of heat-affected powder, phase identification is a viable technique for distinguishing between virgin and used 304L powder.The results from this study show that 304L will change as it is reused in SLM. Even though the degree of its reusability has yet to be determined, changes in the properties of recycled powder can now be better understood with knowledge of the mechanisms responsible for powder degradation. This investigation shows that not only the size distribution of the unconsolidated powder will be altered with continual reuse, but also its chemistry due to the formation of oxides.

This investigation shows that not only the size distribution of the unconsolidated powder will be altered with continual reuse, but also its chemistry due to the formation of oxides. The effects of such changes in powder properties on part quality will be evaluated in the future to assess the reusability of 304L in SLM.AcknowledgementsThis work was funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.References[1] A.T. Sutton, C.S. Kriewall, M.C. Leu, J.W.

ARACTERIZATION OF HIGH ALLOY STEEL PRODUCED VIA ELECTRON BEAM MELTINGDenis Cormier+, Ola Harrysson+, and Harvey West+,++ +Industrial Engineering ++Materials Science and Engineering North Carolina State University Raleigh, NC 27695-7906 Reviewed, accepted August 19, 2003 AbstractElectron Beam Melting (EBM) is a direct-to-metal freeform fabrication technique in which a 4 kW electron beam is used to melt metal powder in a layer-wise fashion. As this process is relatively new, there have not yet been any independently published studies of the high alloy steel microstructural properties. This paper describes the EBM process and presents results of microstructural analyses on H13 tool steel processed via EBM.IntroductionElectron Beam Melting (EBM) is a relatively new direct-metal freeform fabrication process that has been commercialized by Arcam (www.arcam.com). Figure 1 illustrates the basic elements of an Arcam S12 EBM machine.

Figure 1 illustrates the basic elements of an Arcam S12 EBM machine. Typically, a 0.1 mm thick layer of metal powder is first spread across the build platform. A 4 kW electron beam gun then preheats the layer using a relatively low beam current and a relatively high scan speed. This preheating step serves two purposes. First, it lightly sinters the metal powder in order to hold it in place during subsequent melting at higher beam powers. Second, it imparts heat to the part that helps reduce the thermal gradient between the melted layer and the rest of the part. By maintaining a more consistent overall part temperature, built- in residual stresses are reduced. Following this preheating step, the layer is melted by increasing the beam power and/or decreasing the scan speed. The build platform is lowered by an amount equal to one layer thickness, a new layer of metal powder is spread, and the process is repeated until the part is complete.

The build platform is lowered by an amount equal to one layer thickness, a new layer of metal powder is spread, and the process is repeated until the part is complete. Upon completion of the build process, the part is removed from the build chamber. The loose powder supporting the part is removed, and the part is air cooled to room temperature. Bead blasting with silicon carbide, aluminum oxide or other appropriate blast media removes any powder clinging to the part surface. At this point, the as-processed part has a textured surface that resembles a sand casting. If necessary, finish machining of critical part features can be performed.One advantageous feature of the EBM process is that the loose powder supports all downward-facing surfaces, thus highly complex geometric shapes such as conformal cooling channels are easily built without any need for complex 5-axis programming for collision avoidance.

Taminger [1] observed that electron beam melting of metals is very energy efficient and is particularly well suited for processing highly reflective metals such as aluminum that can be difficult to melt with lasers. Lastly, as the entire process takes place in a vacuum, a certain amount of material purification takes place due to the removal of gasses such as those that cause porosity in metal castings.548E-beam gunPowder reservoirPowder spreaderBuild PlatformFigure 1 - Layout of Arcam EBM S-12 MachineDuring the EBM build process, each slice is divided into two regions – contours and squares. The example in Figure 2 only shows a single contour, however, the user can specify that multiple offset contours be built. The area within a layer that remains after the area of the contours is melted is the bulk area to be melted.

The area within a layer that remains after the area of the contours is melted is the bulk area to be melted. Rather than melting all of this area in a left-to-right or spiral scanning pattern, the slicing software subdivides this area into a series of sub-regions called “squares”. In order to evenly distribute heat and to prevent thermal distortions, these squares are melted in a random order that changes from one layer to the next. As is illustrated in Figure 2, squares that are adjacent to a contour are cropped to the shape of the contour. The size of the squares is typically 10 - 15 mm.At present, there do not appear to have been any independently published studies of the Arcam EBM process. The company states that a wide variety of metal powders can be processed with the process, although Arcam initially chose to concentrate on the use of H13 tool steel alloys for tooling applications.

The company states that a wide variety of metal powders can be processed with the process, although Arcam initially chose to concentrate on the use of H13 tool steel alloys for tooling applications. A titanium alloy (Ti6Al4V) has also been developed, and the company is now working on developing parameters for other high- performance alloys. Taminger [1] describes encouraging experiments involving e-beam freeform fabrication of 2219 aluminum alloy in wire form, and Matz [2] also describes unique material properties of a 718 nickel-based superalloy that was fabricated using a similar wire-feed e-beam freeform fabrication process. These studies provide strong evidence that e-beam freeform fabrication is well suited for a wide variety of metal and metal composite materials.

These studies provide strong evidence that e-beam freeform fabrication is well suited for a wide variety of metal and metal composite materials. The remainder of this paper discusses preliminary results of tests conducted using Arcam’s EBM S12 machine to process H13 tool steel powder.Figure 2 – Illustration of Contours and Squares for a Tensile Test Specimen549Figure 3 – A Set of Five Subsize Tensile Test Specimens Prior To Powder RemovalExperimental ProcedureIn order to begin studying the microstructure and properties of H13 steel produced via EBM, several sets of standard 4-inch sub-size tensile test specimens were fabricated in accordance with ASTM E8 [3] (see Figure 3). Parts were built using the default layer thickness of 0.1 mm. Each layer was preheated 10 times with a scan speed of 10,000 mm/sec and a beam current that was progressively increased from 2 mA on the first scan to 20 mA on the 10th scan.

Each layer was preheated 10 times with a scan speed of 10,000 mm/sec and a beam current that was progressively increased from 2 mA on the first scan to 20 mA on the 10th scan. Note that at a scan speed of 10,000 mm/sec, the entire preheating cycle takes a matter of seconds. Following preheating, each layer was melted. A total of two contours were first melted using a beam speed of 100 mm/sec and a beam current up to 20 mA. The squares were then melted at a speed of 500 mm/sec and a beam current up to 20 mA. All processing parameters such as scan speed and beam current can be specified by the user. As is to be expected, the choice of parameter values considerably influences build time and material properties.In total, three jobs consisting of five specimens each were fabricated using identical build parameters as described above. The first two sets of specimens were removed from the S12 machine immediately upon completion of the job and were rapidly air-cooled.

The first two sets of specimens were removed from the S12 machine immediately upon completion of the job and were rapidly air-cooled. As this is an air-hardening steel, it was hypothesized that rapid air cooling would result in the formation of martensite. The second of these two sets was then annealed in a furnace by heating it to a temperature of 900ºC and then allowing it to cool to room temperature at a rate of 22ºC/hr. The third set of specimens was allowed to slowly cool to room temperature in the build chamber under vacuum in order to determine whether or not the properties would be affected by the slower cooling rate (without a subsequent anneal).550Prior to preparation of the specimens for microstructural examination, the hardness of the slowly cooled set of specimens was tested. A set of measurements from different samples produced a range of hardness values from 46.0 to 47.3 HRC.

A set of measurements from different samples produced a range of hardness values from 46.0 to 47.3 HRC. This suggested that cooling in the build chamber under vacuum was not slow enough to prevent the formation of this group were not considered for martensite. Consequently, samples from microstructural examination.One specimen from each remaining group was prepared for microstructural examination. Transverse and longitudinal sections of solid bars were obtained using a water-cooled cutoff wheel and mounted in Bakelite. The surfaces to be viewed were ground on progressively finer grits of silicon carbide paper, rough polished using 6-micron diamond paste, and fine polished with 1 micron and 0.3 micron alumina suspensions. The samples were etched with a 4% solution of nitric acid in methanol. An Olympus inverted stage metallurgical microscope equipped with digital camera was used for optical examination.

The samples were etched with a 4% solution of nitric acid in methanol. An Olympus inverted stage metallurgical microscope equipped with digital camera was used for optical examination. The microstructure was also viewed using a Hitachi S-3200N Environmental Scanning Electron Microscope (SEM) equipped with an Oxford energy dispersive X-ray spectrometer capable of light element detection.For the chemical analysis, samples of the H13 powder as well as a section from a fabricated bar were submitted to NSL Analytical Services for chemical analysis. X-ray fluorescence spectroscopy was used to determine the levels of cobalt, chromium, manganese, molybdenum, nickel, silicon, vanadium, and tungsten.

X-ray fluorescence spectroscopy was used to determine the levels of cobalt, chromium, manganese, molybdenum, nickel, silicon, vanadium, and tungsten. The amount of phosphorus present was measured with optical emission spectroscopy, and the amounts of carbon and sulfur were determined via LECO furnace.Results and DiscussionChemical AnalysisThe results of the chemical composition analysis for the as-processed H13 steel are presented in Table 1. ASTM A681 upper and lower limits [4] for H13 steel are included in the table for reference.

ASTM A681 upper and lower limits [4] for H13 steel are included in the table for reference. Arcam considers its powders to be proprietary at this time, hence composition of the raw powder cannot be included in the table.Table 1- Chemical Composition of as-processed H13 Steel Versus ASTM A681 SpecificationsElementCarbon Manganese Phosphorus Sulfur Silicon Chromium Vanadium MolybdenumMinimum Allowable % 0.32 0.20 0.000 0.000 0.80 4.75 0.80 1.10Measured %0.37 <0.1 0.011 0.005 1.02 4.99 1.05 1.65Maximum Allowable % 0.45 0.60 0.030 0.030 1.25 5.50 1.20 1.75551With the exception of manganese, all alloying elements fall within ASTM limits for H13 steel. One of the interesting challenges with e-beam melting of metal alloys lies in the fact that the process typically takes place in vacuum. Ordinarily, the vaporization of lower melting point alloying elements would be a significant concern.

Ordinarily, the vaporization of lower melting point alloying elements would be a significant concern. With EBM and similar processes, however, solidification of the melt pool takes place so rapidly that the magnitude of vaporization is quite limited. The degree of vaporization is obviously a function of the processing parameters, and future experiments are planned to help obtain a better understanding of how rapidly the metal solidifies under a given set of conditions.HardnessThe hardness of both air-cooled and annealed specimens was measured. The air-cooled specimens had hardness values ranging from 48.0 to 50.0 HRC. This provided initial confirmation of the hypothesis that the as-processed material would consist of martensite. As expected, the specimens that were subsequently annealed had HRC values under 20.Microstructural AnalysisAs described previously, samples of both air-hardened and annealed specimens were prepared for inspection via optical microscopy and SEM.

Figure 4 shows a micrograph of air-cooled steel produced via the EBM process. For reference, a 500 µm scale is shown. Recall that a layer thickness of 100 µm was used to fabricate this part. As expected, the structure is 100% martensite. It is apparent that the part is virtually 100% dense. Interlayer fusion also appears to be complete.500 µmFigure 4 - Microstructure of Air-Cooled H13 Specimen552Figure 5 - Optical Micrograph Showing Local InhomogeneitiesFrom a metallurgical perspective, a set of interesting localized non-homogeneities were observed at the boundary between the melted contours and interior square regions (see Figure 5). These non-homogenous regions are not visible prior to etching. In an attempt to characterize this non-homogeneity, an X-ray line scan across the region shown in Figure 6 was performed to detect variations in the amounts of C, Si, Mo, V, Cr, and Fe present.

In an attempt to characterize this non-homogeneity, an X-ray line scan across the region shown in Figure 6 was performed to detect variations in the amounts of C, Si, Mo, V, Cr, and Fe present. As indicated in the spectra to the right of the figure, the chemical composition within the non-homogeneity is essentially identical to what is found in the surrounding matrix. The specimen was immersed in liquid nitrogen to determine if the non- homogeneity might be retained austenite. Unfortunately, this treatment had no effect on the non-homogeneity. Lastly, a Micromet microhardness tester was used to measure hardness within the non-homogeneity. Microhardness testing indicated that the hardness within the non-homogeneity is identical to that of the surrounding martensite.

Microhardness testing indicated that the hardness within the non-homogeneity is identical to that of the surrounding martensite. Work is continuing on the characterization of these localized non-homogeneities that are only observed after etching.553Figure 6 – SEM Photograph Showing Line Scan and Spectra Analysis Across Inhomogeneity554Figure 7 - Optical Micrograph of Intralayer Shrinkage CrackAlthough the metallurgical properties were generally quite good under the processing parameters that were used, a small number of isolated shrinkage cracks that were contained within specific layers could be observed. One such crack is shown in Figure 7. As the data presented in this paper is specific to one single set of processing parameters, future experiments will be necessary to determine the effect of processing parameters on microstructure and defects.Figure 8 shows a micrograph of an annealed specimen taken at 1000X magnification.

The microstructure is characteristic of spheroidize-annealed steel. The carbides are distributed both along the grain boundaries and within the grains.Figure 9 shows an SEM photomicrograph taken at 30,000X magnification that shows the carbide morphology. As shown in Figure 10, an X-ray spot scan of an individual carbide particle (red) was obtained and compared to the X-ray spectrum of the bulk material (black line). Note the increased amounts of C, Mo, Cr, and V and decreased amount of iron present in the red spectrum.555Figure 8 – Optical Micrograph of Spheroidize-Annealed Microstructure (1000X)Figure 9 - SEM Photograph of Free Carbides In Annealed Specimen (30,000X)556Figure 10 – X-ray Spectra of Bulk Material and Carbide In Annealed SpecimenConclusionsThis paper has presented a microstructural analysis of H13 tool steel produced via the EBM process.

As the processing conditions will significantly affect metallurgical properties, the decision was made to conduct this initial study using default processing conditions. Although parameters can be “tuned” for each build according to the part’s geometry, the metallurgical properties obtained using the default parameter values were quite good. The parts exhibited full interlayer bonding with virtually no porosity. The as- processed material was martensite having a hardness of 48-50 HRC. The spheroidize- annealed material also appeared much as expected, with carbides dispersed along the grain boundaries and within the grains themselves.An interesting observation was the presence of localized non-homogeneities along the boundary between the part’s exterior contours and interior squares. The chemical composition and hardness of these regions is virtually identical to that of the surrounding martensite.

The chemical composition and hardness of these regions is virtually identical to that of the surrounding martensite. Work continues to further characterize these regions.As one would expect, the selection of processing conditions has a significant influence on build speed as well as mechanical properties. Generally speaking, an increase in build speed is accompanied by a decrease in mechanical properties. The next phase of experimentation (just underway) is to begin studying the relationships between mechanical properties and parameters such as the number of contours, the square sizes, the beam currents (for contours and squares), the beam speed (for contours and squares), the layer thickness, etc. These experiments will help map out the tradeoffs between build speed and metallurgical properties so that the processing parameters can be selected that best meet the needs of the application.557References[1] Taminger, K. and Hafley, R.

and Hafley, R. (2002) “Characterization of 2219 Aluminum Produced by Electron Beam Freeform Fabrication” 2002 Solid Freeform Fabrication Symposium Proceedings, pp. 482 - 489[2] Matz, J. and Eagar, T. (2002) “Carbide Formation in Alloy 718 during Electron Beam Solid Freeform Fabrication” Metallurgical and Materials Transactions A, vol. 33A, pp. 2559 – 2567.[3] Standard Test Methods for Tension Testing of Metallic Materials, ASTM E8-01.[4] Standard Specification for Tool Steels Alloy, ASTM A681-94.558.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCHARACTERIZATION OF HIGH-DEPOSITION POLYMER EXTRUSION IN HYBRID MANUFACTURINGJason Weaver*, Jason Jones†Dept. Manufacturing Engineering Technology, Brigham Young University, Provo, UT 84602 †Hybrid Manufacturing Technologies, McKinney, UT 75069AbstractHybrid manufacturing processes that include additive and subtractive operations have unlocked many of the design limitations that were not previously available. Additive processes allow increased design flexibility, customization, and complexity. Subtractive processes enable higher production speeds and improved accuracy and surface finish. This paper describes a hybrid system that combines a computer numerically controlled (CNC) milling machine with a high deposition rate polymer extruder to create artifacts using hybrid additive/subtractive processes.

Future research plans are described, including possible applications for this system in multimaterial and multi-process manufacturing.IntroductionAdditive manufacturing processes offer many advantages over conventional subtractive processes. Foremost among these are the expanded design space enabled by layer-by-layer deposition and the low barriers to mass customization and increased complexity. However, additive manufacturing processes also carry well known shortcomings, including longer build times, inferior surface finish, and less control of dimensional accuracy. Subtractive processes offer the opposite—more restrictions on design complexity and customization, but higher output and better control of surface finish and tolerances.

Subtractive processes offer the opposite—more restrictions on design complexity and customization, but higher output and better control of surface finish and tolerances. Many methods have been explored to combine the best of both additive and subtractive processes, an idea typically called hybrid manufacturing.One approach that has been recently explored is combining a metal powder directed energy deposition (DED) process with a multi-axis CNC milling process. This can be a custom machine that incorporates both processes, or it can be an add-on module that adds a second process to a commercially available system. Hybrid Manufacturing Technologies offers several such tools, such as the AMBIT (Figure 1). The AMBIT is an interchangeable tool with a CAT 40 head that can be mounted and controlled by a CNC mill.287Figure 1.

The AMBIT is an interchangeable tool with a CAT 40 head that can be mounted and controlled by a CNC mill.287Figure 1. AMBIT metal DED toolAnother technology that is relevant to this paper is that of room-scale polymer extrusion systems, such as big area additive manufacturing (BAAM) (Figure 2). In these systems, large nozzles, efficient heating and shearing subsystems, and the use of polymer pellets instead of filament as feedstock enable high speed fabrication with thick layer heights and large part footprints.Figure 2. "Strati" 3D-printed car, manufactured in part through BAAM processAmbit PE-1These two concepts have been combined in Hybrid Manufacturing Technologies’ AMBIT PE-1 polymer extrusion tool (Figure 3). A prototype version of this system has been installed on a Haas toolroom mill at Brigham Young University, where its capabilities and possible applications will be investigated.

A prototype version of this system has been installed on a Haas toolroom mill at Brigham Young University, where its capabilities and possible applications will be investigated. The AMBIT PE-1 uses the same CAT 40 head as the other AMBIT tools, and includes a miniaturized auger driven by the CNC machine’s spindle. This auger288helps to shear and blend the polymer pellets as they are melted by a heater in the tool. External connections are needed to the temperature control and the pellet hopper. The tool can print in common polymers like ABS up to 9 kg/hr, and can be switched out for subtractive CNC tools or inspection tools to enable true hybrid manufacturing in a single commercially available set up. Because the same pellets are used as in injection molding machines and other well-established processes, a wide variety of possible materials are readily available for cheaper than the equivalent filaments.

Figure 4 shows the prototype head installed on the Haas CNC mill at BYU.Figure 3. Hybrid Manufacturing Technologies AMBIT PE-1 polymer extruder for CNCsFigure 4. AMBIT PE-1 prototype installed on a HAAS CNC mill at BYUProposed Research Road MapResearch regarding this hybrid manufacturing process will occur in four general stages. During the first year, to occur in fall 2018, our students will complete the set up and characterization of the AMBIT PE-1 tool. In particular, we will determine optimum settings for the default material—ABS with chopped carbon fiber—and consider which parameters are tuned similar to calibration of FFF machines, which behave more like BAAM, and which may behave289in unique ways in this application.

We will also print test specimens in CF-ABS to compare material properties to similar materials printed on FFF or injection molded.During the second stage, after the system is performing reliably, we will investigate the viability of using this system for both additive and subtractive processes. In metal DED hybrid systems, the additively produced material is welded securely to the build plate, allowing subsequent milling and other subtractive processes without refixturing the part. However, polymer systems typically secure the printed material to the bed by using similar materials (e.g., printing ABS onto an ABS sheet), softening the bottom layers on a heated bed, using adhesives like glue or tape, or allowing the material to seep into small perforations on the build plate.

These methods are sufficient to prevent the part from coming loose during fabrication, but may not be secure enough to secure the part against the torques and horizontal pressures experienced during milling operations. We will investigate and test variations of these existing solutions, as well as possible novel solutions like print-in-place clamps and mechanical interlocks (like tongue-in-groove) with the bed.The third stage, which can be run in parallel with the second, is to characterize the system and adapt it for use with high performance and high temperature polymers. These polymers, such as PEEK, PEK, PTFE, PPS, PEI, and PC, are in high demand for engineering applications where high strength and high temperature survivability is needed. However, these materials are difficult and expensive to find in filament form, and most FFF machines are not set up for such extreme temperatures.

However, these materials are difficult and expensive to find in filament form, and most FFF machines are not set up for such extreme temperatures. With the AMBIT PE-1’s use of pellets, these materials can more easily be used in additive designs.The fourth planned stage of research is to consider possible applications for this technology, especially opportunities for using the process in multimaterial and multi-process designs. Possible hybrid combinations where this tool could be advantageous include cladding high performance materials over less expensive bulk materials for tools and molds, printing polymer on metal substrates, printing metal on high-temperature polymer, continuous polymer functional gradient materials (FGMs), and even embedded electromechanical sensors and motors.

The eventual goal would be the incorporation of additive manufacturing into true CNC automated manufacturing, with single-program multimaterial hybrid process planning and automatic ejection of the finished part.290.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCharacterization of Impact Toughness of 304L Stainless Steel Fabricated through Laser Powder Bed Fusion ProcessSreekar Karnati1, Atoosa Khiabani1, Aaron Flood1, Frank Liou1, Joseph W. Newkirk21Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409 2Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractIn this research, the impact toughness of powder bed based additively manufactured 304L stainless steel was investigated. Charpy specimens were built in vertical, horizontal and inclined (45o) orientations to investigate the variation in toughness with build direction. These specimens were tested in as-built and machined conditions.

These specimens were tested in as-built and machined conditions. A significant difference in toughness was observed with varying build directions. The lowest toughness values were recorded when the notch was oriented in line with the interlayer boundary. The highest toughness was recorded when the notch was perpendicular to the interlayer boundary. A significant scatter in toughness values was also observed. The variation and distribution among the toughness values were modeled by performing 3-parameter Weibull fits. The performance and variation of the additively manufactured 304L were also compared with the toughness values of wrought 304 stainless. The additively manufactured material was observed to be significantly less tough and more variant in comparison to wrought material.IntroductionAdditive manufacturing (AM) in recent times has risen from being a solely prototyping technology to a full-scale manufacturing technology.

Especially for metals, various commercial solutions based on blown powder and powder bed methodologies have been developed to address the demands of the industry. Multiple vendors such as Renishaw, Concept Laser, SLM Solutions, EOS, etc. have developed robust solutions capable of fabricating complex geometries from multiple alloys. Popular alloys such as SS304, Inconel alloys, titanium alloys, precipitation hardenable steels etc. are currently being incorporated into these machines. Despite using similar chemistries as feedstocks, in comparison to conventional processes, parts with drastically different properties are obtained through AM. Such strong differences in material behavior currently limit wide scale incorporation of AM into various industries.Multiple studies have been performed on comparing the properties of AM material with conventionally fabricated counterparts.

The strength properties of AM material have been reported to be similar if not better than their counterparts. However, the properties of AM material were often noticed to be anisotropic. Conclusions from these studies indicate strong direction dependence in mechanical performance [1–13]. The directional heat transfer and rapid solidification in AM process result in a very orderly and repetitive constitution. This type of structure within the AM material is expected produce direction dependent performance.©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1336Yang et. al. reported anisotropic tensile performance in Ti6Al4V fabricated through laser powder bed fusion. The directional dependence was attributed to the crystallographic constitution of the material. This anisotropy was also observed to change with input energy density [12]. Chlebus et. al.

The directional dependence was attributed to the crystallographic constitution of the material. This anisotropy was also observed to change with input energy density [12]. Chlebus et. al. also reported anisotropy in powder bed fabricated Ti6Al7Nb. While strengths higher than conventionally fabricated material were reported, lower ductility and directionally varying performance were also noted [11]. Similar observations in AM fabricated stainless steels were also reported by Suryawanshi et. al.[8], Deev et. al.[9], and Zhukov et. al. [1]. The identification and characterization of such differences are vital to designers. Knowledge of anisotropy is critical for designing structural components. In order to achieve safe and reliable performance, appropriate compensations are necessary to account for the variable properties.In the current work, the variation of impact toughness with build direction was characterized for powder bed fabricated 304L stainless steel.

Impact testing of specimens in as- built and machined conditions was performed. The performance of AM stainless steel was compared against that of wrought material. The scatter in the toughness performance of both materials was also modeled.Experimental setupThe AM material in the current study was fabricated using the Renishaw AM250 machine located at Missouri S&T. The parameters used to fabricate the specimens for the study were optimized for attaining maximum density and high quality downward facing surfaces. The Renishaw machine follows a point-by-point exposure method as opposed to a continuous laser scan which is typical for most laser based powder bed machines. During this point by point exposure, a power of 200W, a point spacing of 65 micron, an exposure time of 75 micro seconds and a hatch spacing of 85 micron were used. The chemistry of the powder used for AM fabrication is listed in Table 1.

The chemistry of the powder used for AM fabrication is listed in Table 1. For the comparative study, a 7/16” wrought 304 stainless steel bar stock (rolled and annealed) was also analyzed. The composition of the wrought material is listed in Table 2.Element Wt. %Table 1. The chemical composition of 304L stainless steel powder C 0.018Cr Fe Mn 18.4 < 0.1 Bal 1.4CuN 0.06Ni 9.8O 0.02P 0.012S 0.005Table 2. The chemical composition of wrought 304L stainless steel bar stockElementWt. %C0.03Cr 18- 20CuFe MnBal2N0.01Ni 8- 12OP0.045S0.03The Charpy specimens were prepared to a size of 55mmx10mmx10mm. The 2mm deep “V” notch was machined using a standard broach. Two sets of AM specimens were built for Charpy testing. These specimens were built in vertical, horizontal and 45o orientations. A screenshot of the build layout is shown in Figure 1. The two sets of specimens were prepared for testing in as-built and machined conditions ().

A screenshot of the build layout is shown in Figure 1. The two sets of specimens were prepared for testing in as-built and machined conditions (). The specimens were built in the same layout for©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1337Si 0.6Si1both of the surface conditions. The specimens for testing in the machined state were built oversized to account for the machining compensation. For comparison, specimens made from wrought stainless steel were also tested. The wrought raw stock was machined to dimensions and notched for testing.

For comparison, specimens made from wrought stainless steel were also tested. The wrought raw stock was machined to dimensions and notched for testing. The total number of specimens tested in each case are shown in Table 3.Figure 1: Build layout for Charpy specimens in different orientations, XY directions along the plate Z direction same as the build direction (out of the paper).Table 3: Total number of specimens tested for impact toughness.MaterialOrientation# of specimens023SLM, As-built45209025024SLM, Machined45219024WroughtNotch perpendicular to the rolling direction30Results and DiscussionsThe densities of the machined AM specimens measured by Archimedes method in distilled water, categorized by build orientation are shown in Figure 2. While there was no statistically significant difference in densities, the variance and median values for each orientation were observed to be different.

While there was no statistically significant difference in densities, the variance and median values for each orientation were observed to be different. The variance among the specimens built along the build direction was observed to be the highest. The median density of the 45 degrees specimens was the highest and vertical specimens was the lowest. The large density variance in vertical specimens is suspected to be due to the difference in probability of formation of lack of fusion defects. While the parameters were optimized for maximum density, several environmental factors (humidity,©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1338oxygen entry during powder feed, inconsistency in wiper setup etc.) can influence the probability of pore formation and in extension the part density.

The specimens built in the vertical orientation require the highest number of layers for completion of fabrication and hence can contain large variance in density. The density is also influenced by the build location, lens transmission deterioration (accumulation of condensed metal, ejecta, fines etc.).Interval Plot of Density (gm/cm^3) 95% CI for the Mean7.967.94) 3 ^ m c / m g (7.927.90y t i s n e D7.887.867.847.82045 Orientation (Degrees)90Individual standard deviations are used to calculate the intervals.Figure 2: Density values of machined specimens built in 0, 45, and 90-degree orientations with respect to the substrate.

‘●’ is the mean and ‘▼’is the median, the connecting line joins the medians, the error bars span a range of 2 standard deviations.250200) J (s s e n h g u o T150100500Orientation State045 AsBuilt90045 Machined90045 Wrought90Figure 3: Box plot of toughness values measured from AM specimens in as-built and machined states and wrought specimens.©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1339For specimens tested in both as-built and machined state, the toughness values of the specimens built vertically were substantially lower than those of horizontal and 45-degree specimens. The toughness values of 45-degree specimens were observed to be slightly lower when tested in the machined state. However, the difference was observed to be insignificant when tested in the as-built state.

However, the difference was observed to be insignificant when tested in the as-built state. The same can be seen from the box plots shown in Figure 3. Overall the toughness values of specimens in as-built conditions were observed to be substantially higher than those of machined condition. This observation is counter-intuitive, the poor surface quality is expected to result in early failure and lower toughness. This difference in toughness is suspected to be an artifact of the differences in the testing of as-built and machined specimens.While the as-built specimens were designed to be built to size, they are consistently oversized from the specification (by approx. 0.1mm). Although this difference is small, it is still consistent. This small difference in material volume could add to the toughness value of the as- built specimen. The surface quality between these specimens is also dramatically different.

This small difference in material volume could add to the toughness value of the as- built specimen. The surface quality between these specimens is also dramatically different. The as-built specimens have a rough finish which is an artifact of the build process, whereas the machined specimens have a smooth finish. The difference in friction can add to the toughness of the as-built specimens. Also noticeably different plastic deformation was observed in the regions where the specimens come in contact with the pins (See Figure 4). The as-built specimens underwent more deformation in comparison to the machined specimens leading to extra energy absorption and high toughness measurements.Figure 4: Difference in plastic deformation from indentation of the pins during Charpy testing (red arrows).

The threshold value is an estimate of the value under which the probability of the event occurring is zero. From a designer’s perspective, the threshold value is the most critical input. Basing the threshold value as the failure criteria, maximum life and performance can be reliably expected from a component. The Weibull fits of data from machined and wrought specimens are shown in Figure 5. The values of the Weibull parameters are shown in Figure 6 .Similar to the conclusions from the box plots in Figure 3, Weibull representation also indicates the variation in vertical samples was substantially higher than the remaining orientations (see Figure 5). The lowest toughness value in the vertical orientation is substantially smaller than those in horizontal and inclined orientation. The performance from the inclined orientation is slightly inferior to that of the horizontal orientation. However, the scatter in the performance is similar.

The performance from the inclined orientation is slightly inferior to that of the horizontal orientation. However, the scatter in the performance is similar. The drastic differences in the threshold values indicate the need for considering material anisotropy during component design.Figure 5: Three parameter Weibull fits of toughness data for powder bed (left) and wrought (right) 304 stainless steel.) J (s s e n h g u o T200 150 100 50 0SLM, 0 degreesSLM, 45 degreesSLM, 90 degreesWrought3210s u u d o mll lu b e WiThresholdCharacteristic strength Weibull modulusFigure 6: Summary of Weibull parameters from the fits shown in Figure 5.©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1341The threshold values of AM material were substantially lower than those of wrought material.

While the Weibull modulus of wrought material was lower than the Weibull modulus of horizontal AM specimens, the difference between the threshold and characteristic strength were substantially different for the horizontal AM specimens. The Weibull modulus values were observed to decrease with orientation for the AM specimens, 0 degrees was the highest and 90 degrees was the lowest. Overall, the AM material performance is significantly inferior and scattered in comparison to wrought material.In order to understand the anisotropy and the inferior performance of AM material, the fracture surfaces of the broken Charpy specimens were investigated. The fracture surfaces of the specimens with lowest, median and highest toughness values were imaged using ASPEX SEM.

The fracture surfaces of the specimens with lowest, median and highest toughness values were imaged using ASPEX SEM. The complete stitched images of the fracture surfaces are shown in Figure 7.Figure 7: Stitched images of fracture surfaces of specimens with lowest (left column), median (middle column) and highest (right column) values of toughness in vertical (top row), 45 degrees (middle row) and horizontal (bottom row) orientations.©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1342The toughness values of the specimens are in agreement with the deformation of the cross- section. The lowest performers have a near zero deformation of the cross section, suggesting a brittle failure. Whereas the best performers have significant deformations indicating ductile failure.

The lowest performers have a near zero deformation of the cross section, suggesting a brittle failure. Whereas the best performers have significant deformations indicating ductile failure. The vertical specimen with the lowest toughness value in comparison to the horizontal specimens with both lowest and highest toughness value has a near square cross-section suggesting minimal amounts of deformation. The lack of deformation and low toughness value could be indicative of a failure mechanism whose sensitivity varies with build orientation. While dimples, representative of ductile failure, were present across the fracture surface, features in the shape of smoothly rounded pits were also observed (see Figure 8). The shape of these pits was observed to vary with orientation. From the shape and size of these features, these pits appear to be track boundaries. This could imply that the weakest link in this AM material could be track boundaries.

From the shape and size of these features, these pits appear to be track boundaries. This could imply that the weakest link in this AM material could be track boundaries. Furthermore, the large variance and low toughness of vertical specimens could mean that the inter-layer track boundary is the weakest link in the AM material.Figure 8: Close up images of fracture surfaces of horizontal, vertical and 45 degrees specimens with median toughness values.Macro images of fracture surfaces from broken Charpy specimens are shown in Figure 9. From the figure, it can again be observed that the fracture surfaces possess an atypical pattern for ductile materials. Features such as steps, grooves, and ridges can be seen on the fracture surface. Especially, the fracture surface of the vertically built specimens indicates sharp changes in the direction of crack propagation. The step-like features are expected from 90 degree turns in the direction of crack propagation.

The step-like features are expected from 90 degree turns in the direction of crack propagation. The step-like features could be a result of the joining of cracks initiated in different layers.If the “V” notch is in between layers (vertical samples), step-like features were observed. If the notch is perpendicular/inclined to the interlayer boundary, grooves/ ridges have been observed. Considering the location of the notch and the patterns on the fracture surfaces, the propagation of the crack is expected to be exclusively along the interlayer boundary.©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1343Figure 9: Fracture surfaces of broken Charpy specimens built in different build orientations.To confirm the above claims, cross sections of broken specimens were cut, ground, polished and etched to reveal the microstructure (shown in Figure 10).

The microstructure along the fracture surface substantiates the above claims. The fracture surface was seen to be along the track/interlayer boundary. The step-like features were formed due to the changes in the direction of crack propagation from one interlayer boundary to another along the track boundaries in between those two sets of layers. The difference in build orientations causes a difference in location of the notch with respect to the interlayer boundaries. This makes the vertical specimens the most susceptible and the horizontal specimens the most resistant to failure. This also explains the difference in the type of patterns observed on the fracture surface.Figure 10: Microstructure along the fracture surface on a plane normal to the notch on broken specimensThe low performance of the AM material despite high densities and the atypical crack propagation along the track boundaries implies the presence of a critical flaw along the track boundaries.

The ordered and repetitive constitution of AM material also orders the said flaw throughout the fabricated component. Upon loading, the performance of the AM material is dependent on the access of the crack to this flaw distribution. While the presence of the flaw was not identifiable under low magnification imaging, further investigation in required to explore this failure mechanism.©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1344ConclusionsThe impact performance of AM 304L stainless steel material fabricated using laser powder bed fusion was studied against that of wrought 304L stainless steel.While the specimens had high densities, the performance of the AM material was inferior to that of wrought material.A substantial amount of anisotropy in impact toughness was observed with varying build direction.

The vertically built specimens exhibited the lowest toughness performance. • The scatter in the toughness performance of the vertically built specimens was also substantially higher than that of the other two build orientations.The fracture surfaces of the broken specimens indicate an atypical failure pattern for ductile materials.The crack propagation was exclusively along the track boundary. The varying access of the crack to the interlayer track boundary was identified to be the source of anisotropy.AcknowledgmentsThis work has been funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy.

DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes. The supports from National Science Foundation Grants CMMI-1625736, and the Intelligent Systems Center (ISC) at Missouri S&T are greatly appreciated.Furthermore, the authors would like to acknowledge Kyle Stagner, Ian Wille, and BrianBullock for their aid in this research.References[1] A. Zhukov, A. Deev, P. Kuznetsov, Effect of Alloying on the 316L and 321 Steels Samples Obtained by Selective Laser Melting, Phys. Procedia. 89 (2017) 172–178. doi:10.1016/J.PHPRO.2017.08.010.[2] Y.M. Wang, T. Voisin, J.T. McKeown, J. Ye, N.P. Calta, Z. Li, Z. Zeng, Y. Zhang, W.Chen, T.T. Roehling, R.T.

ARACTERIZATION OF IN-SITU CONDUCTIVE PASTE EXTRUSION ON POLYJET SUBSTRATESK Blake Perez and Dr. Christopher B. WilliamsDesign, Research, and Education for Additive Manufacturing Systems Lab Department of Mechanical Engineering, Virginia Tech, Blacksburg, VA 24061ABSTRACTThe integration of Direct Write technologies into Additive Manufacturing system enables the in-situ deposition of conductive traces during part printing, and thus the creation of parts with embedded electronics. In this paper, the authors detail their research of integrating an extrusion-based direct write system into a PolyJet material jetting system to create multi- material products with structurally integrated, functional electronics. An investigation of the dispensing (e.g. orifice diameter, dispense pressure, and toolhead speed), drying (e.g., time and temperature), and substrate parameters (e.g., VeroWhitePlus and TangoBlackPlus) on the geometry of the deposited trace is presented.

Additionally, the adhesive compatibility of the conductive material on both rigid and elastomeric PolyJet substrate surfaces is investigated by measuring wet and dry contact angles.1INTEGRATION OF DIRECT WRITE AND ADDITIVE MANUFACTURINGThe integration of Direct Write (DW) technologies, which can be used to selectively deposit conductive materials, with Additive Manufacturing (AM) processes can enable the realization of mechatronic products that feature embedded actuation, sensing, and power.

Moreover, integration of DW and AM enables the embedding of the electronics within the component’s structure, thus eliminating the need for external wiring and interconnects and providing a means for protecting the electronics from the environment.As seen in Table 1, prior efforts in depositing conductive traces onto additively manufacturing parts has been focused in integrating DW with single-material stereolithography, ultrasonic consolidation, extrusion, and laser sintering AM technologies. However, none incorporate Material Jetting processes.Table 1. Prior Combinations of Direct Write and Additive Manufacturing Technologies Ref.

However, none incorporate Material Jetting processes.Table 1. Prior Combinations of Direct Write and Additive Manufacturing Technologies Ref. [1] [2] [3] [4] [5] [6] [7]Application Interconnects Discreet Electronics Interconnects Batteries Conformal Electronics Antenna & Interconnects InterconnectsDW Technology Micropump Pneumatic Extrusion Pneumatic Extrusion Screw Extrusion Manual Aerosol Jet Aerosol Jet & Micropump Laser SinteringAM Technology Stereolithography Stereolithography Ultrasonic Consolidation Screw Extrusion Stereolithography Fused Filament Fabrication1A comprehensive review of combining direct write and additive manufacturing technologies can be found in [8].

In general, the major challenges in integrating DW + AM include:§ Z-Direction Electrical Interconnects: Depositing multiple and electrically continuousconductive structure in multiple layers.§ Process Integration: Combining different DW & AM technologies in a single process § Material Compatibility: Ensuring that both DW & AM material are compatible in terms ofadhesion, resolution, and thermal processing.In this work, the authors explore the integration of extrusion-based DW technology with the Stratasys PolyJet material jetting AM technology. PolyJet is considered specifically because of its ability to simultaneously process both rigid and elastomeric photopolymers, which allows for the direct fabrication of mechanisms with actuated living-hinge joints [9]. The integration of PolyJet and DW could enable direct manufacture of complex, multi-material components that feature structurally integrated actuation and sensing in a single build.

The integration of PolyJet and DW could enable direct manufacture of complex, multi-material components that feature structurally integrated actuation and sensing in a single build. An overview of the embodied DW/PolyJet integration is provided in Section 2.The goal of this work is to gain an understanding of the compatibility of the materials processed by the two material deposition technologies. Specifically, in this paper, the authors explore how process parameters of both the DW and the PolyJet processes affect the deposited materials. The research is guided by three questions:1. How do tip size, dispense pressure, and tool head speed affect feature width and height?2. How does drying time and temperature affect the geometry of conductive inks on PolyJet substrates?3.

How do tip size, dispense pressure, and tool head speed affect feature width and height?2. How does drying time and temperature affect the geometry of conductive inks on PolyJet substrates?3. What are the adhesion characteristics between the conductive ink and the PolyJet substrates?Regarding Question 1, two experiments are performed to explore both the width and height of the deposited conductive geometries under various dispensing and drying parameters on both VeroWhitePlus (rigid) and TangoBlackPlus (elastomeric) PolyJet substrates (Section 3). Using the determined ideal dispense parameters (for thinnest depositable traces), a second experiment is performed to analyze the height of the dispensed traces and its relationship to drying parameters (Question 2, Section 4). To answer Question 3, the authors investigate contact angle between the PolyJet substrates and conductive ink as a means of determining adhesive compatibility between the two materials (Section 4).

To answer Question 3, the authors investigate contact angle between the PolyJet substrates and conductive ink as a means of determining adhesive compatibility between the two materials (Section 4). Finally, a discussion synthesizing lessons from all three experiments is provided for closure is presented in Section 6.2INTEGRATION OF DIRECT WRITE WITH POLYJET MATERIAL JETTINGAs noted by the summary of literature presented in Table 1, no prior work has been demonstrated in combining DW and material jetting AM technologies. To enable the selective deposition of conductive materials during the PolyJet printing process, a separate deposition machine was realized. The system’s embodiment was performed according to design consideration criteria for combining DW and AM technologies, as presented in [3]. The resultant DW setup is presented in Figure 6.2Figure 6.

The resultant DW setup is presented in Figure 6.2Figure 6. Direct Write setup for PolyJet processThe guiding design goal for the realization of the system was that it needed to be able to be placed onto the PolyJet build tray in order to deposit conductive material directly onto an in-situ part. As such, an extrusion-based DW technology was selected, as it is easily placed into the PolyJet system. Specifically, the infrastructure of extrusion processes is compatible with the PolyJet machine as it does not require a closed environment and can be incorporated in an appropriately-sized positioning system. In addition, extrusion DW systems have flexibility in processing conductive materials of varying viscosities. Specifically, extrusion-based DW is capable of dispensing heavily loaded inks; previous works have used inks with metal loadings in the 60-70% range [4]–[7].

Specifically, extrusion-based DW is capable of dispensing heavily loaded inks; previous works have used inks with metal loadings in the 60-70% range [4]–[7]. Additionally, the 50 µm feature resolution is comparable to the PolyJet process (42 µm); although this resolution value may increase with the use of heavily loaded inks.A Nordson EFD Ultimus V high precision dispenser was selected for the extrusion-based dispensing. The dispenser uses air pressure to dispense materials via a syringe. Positive pressure from 0-100 psi in increments of 0.1 psi can be applied to the syringe to extrude material, and a vacuum can be applied quickly to stop the material flow. Additionally, it has the ability to consistently meter materials from the syringe as the material level changes (nearly full versus nearly empty syringe).

Additionally, it has the ability to consistently meter materials from the syringe as the material level changes (nearly full versus nearly empty syringe). This removes the material metering control loop out of the larger control system, thus allowing for repeatable deposition control.The syringe of the extrusion-system is mounted onto an X-Y frame, which is made from extruded aluminum framing, and is sized to fit on the Connex 350 build tray. Stepper motors are used to drive lead screws for X- and Y-direction translation. A one-inch lead screw pitch and 200 pole stepper motors result in a minimum system resolution of 125 µm per step. Translation is controlled by an Arduino programmed to interpret basic g-code. Dispense commands are also driven by the Arduino and interpreted from the g-code.

Translation is controlled by an Arduino programmed to interpret basic g-code. Dispense commands are also driven by the Arduino and interpreted from the g-code. An overhead camera (not shown in figure) was used with LabVIEW’s vision and motion toolbox to provide image-based orientation alignment between the inserted extrusion platform and the PolyJet build tray.In this work, the extrusion system is used to selectively deposit DuPont 5021 silver-loaded conductive ink. This ink was chosen as its post-processing temperatures were compatible with3the low heat-deflection temperature of printed PolyJet substrates. Because the goal is to encapsulate the directly written traces after deposition with subsequently printed layers, the components cannot be removed from the build tray and treated in an oven.

Because the goal is to encapsulate the directly written traces after deposition with subsequently printed layers, the components cannot be removed from the build tray and treated in an oven. (Removing the components would require them to be replaced on the build tray with microscale precision.) As such, a heat gun is used to force warm air over the samples during the drying process on the build tray.3 EXPERIMENT 1: DW PROCESS PARAMTERS’ EFFECT ON TRACE THICKNESS3.1 ContextDuring an assessment of the design considerations for hybridizing additive manufacturing and direct write technologies, depositable feature size was identified as a key metric for assessing the quality of the printed conductive traces.

The following question guided the search of the DW process parameters that provided the smallest deposited feature: How do tip size, dispense pressure, and tool head speed affect feature width?For the pneumatic extrusion-based dispenser selected, tip size, dispensing pressure, and tool head speeds are the main process parameters that influence the geometry of deposited beads of material [14]. While the review of direct write technology suggests that deposition resolutions from pneumatic solutions can match that of the PolyJet process [8] this assumption must be validated. Beyond the resolution qualification, this experiment aims to assess how different dispensing parameters affect feature width on different PolyJet substrates.During the dispensing process, the material passes through the tip orifice and the extrudate assumes the cross-sectional geometry of the opening. As the material reaches the substrate, it is expected to slump, creating a trace wider than the tip orifice.

As the material reaches the substrate, it is expected to slump, creating a trace wider than the tip orifice. Nonetheless, smaller tip diameters will yield smaller features relative to features from larger tip sizes. However, there is a practical limit to the minimum tip diameter used based on the material and clogging. Heavily loaded inks will clog smaller tip orifices. Because pressure is proportionally related to the volumetric flow rate, it is hypothesized that lower pressures will produce lower flow rates and thus smaller features, although insufficient pressure will not allow material extrusion.Moreover, faster tool head speeds should also produce smaller feature sizes. For a given pressure, the mass flow rate of conductive material exiting the nozzle is constant.

For a given pressure, the mass flow rate of conductive material exiting the nozzle is constant. Equation 1 relates the mass flow rate (𝑄) to the cross-sectional area of the deposited conductive trace (Acs) and the toolhead speed (Vtoolhead).𝑄 = 𝐴!" ∙ 𝑉!""#!!"#If velocity increases, the cross-sectional area will decrease. It is possible that this will cause the beads to become shorter, thinner, or both. The next section discusses the methods used to investigate this research question.3.2 MethodIn designing the experiment to assess the effects of the main factors (tip size, pressure, and tool head speed), it is assumed that there is no significant interaction between them; i.e., each factor should affect Equation 1 independently. However there is no specific proof when considering PolyJet substrates. Therefore, a full factorial analysis is performed with each main4(1)factor combination on both PolyJet substrates VeroWhite+ and TangoBlack+.

Therefore, a full factorial analysis is performed with each main4(1)factor combination on both PolyJet substrates VeroWhite+ and TangoBlack+. The main experimental factors, and their variable ranges are outlined in Table 2.Table 2: Selected main factor variable settingsVariableTip SizeDispensing PressureSubstrateToolhead SpeedUnit[um][kPa][ ][mm/min]Range[400][31.03-41.81][VeroWhite+ TangoBlack+][1000-3000]Rationale Smaller tip sizes will produce finer features although there is a practical limitation where the conductive material will no longer extrude due to nozzle clogging. The selected range is also influenced by commercially available tip sizes for the pneumatic dispenser. Equation 1 demonstrates that lower pressures will decrease the volumetric flow rate and thus lead to smaller features. Preliminary results show that pressures below 31.03 kPa are insufficient for extrusion.

Preliminary results show that pressures below 31.03 kPa are insufficient for extrusion. Pressure settings 37.92 and 41.81 kPa are also tested to observe potential interactions. These are the two main materials for PolyJet printing. VeroWhite+ allows for rigid features, while TangoBlack+ is elastomeric and flexible. Equation 1 demonstrates that faster toolhead speeds should also allow smaller features. 3000 mm/min is selected as an upper limit because the deposition machine is not capable of depositing reliably at faster speeds.Preliminary screening eliminated tip size as a main factor when processing the DuPont 5021 material. It was observed that tips with a diameter smaller than 400 µm would clog repeatedly. At times, the metal particles clogged the nozzle to allow only the solvent portion of the ink to extrude.

It was observed that tips with a diameter smaller than 400 µm would clog repeatedly. At times, the metal particles clogged the nozzle to allow only the solvent portion of the ink to extrude. Therefore, experiments were only conducted with the combinations of the main factors using the 400 µm tip diameter.For dispensing pressure, pressures less than 31.03 MPa were not sufficient to extrude material and those over 41.81 MPa caused material agglomeration on the substrate surface and resulted in many discontinuities. The range of toolhead speeds is determined by the maximum speed at which the X-Y motion system loses its positioning accuracy. While extrusion at faster speeds is possible, it is not done reliably with the current machine embodiment.

While extrusion at faster speeds is possible, it is not done reliably with the current machine embodiment. Beyond this speed barrier, there is expected to also be a physical limitation to toolhead speed based on the rheology of the conductive ink.There are 18 different combinations in total shown in Table 2, and each combination is tested four times. For each factor combination, a 40mm long line is deposited onto one of the two printed substrates. Width data of the resulting deposited line is measured using a HIROX KH-7700 3D digital video microscope.3.3 ResultsThe average bead width data for each main factor combination is shown below in Table 3. The highest pressure setting (44.81 kPa) induced nozzle clogging and resulted in discontinuous lines. As a result, width data is not gathered for that particular setting. For varying pressure, lower pressures exhibited thinner lines. In addition, faster toolhead speeds resulted in thinner lines.

As a result, width data is not gathered for that particular setting. For varying pressure, lower pressures exhibited thinner lines. In addition, faster toolhead speeds resulted in thinner lines. When comparing both substrates, lines written on TangoBlack+ were on average 343 µm thinner than those on VeroWhite+.5An ANOVA analysis of the main factors did not demonstrate any significant interactions, as shown by the interaction plot in Figure 1. The lines in these plots are mostly parallel and represent the different setting for the parameter listed in that row. It is possible that there may be a slight interaction between speeds and dispense pressures as the green and blue lines show some convergence.

It is possible that there may be a slight interaction between speeds and dispense pressures as the green and blue lines show some convergence. Therefore, higher speeds may have more of a dependence on pressure as shown by the blue line (3000 mm/min) that has a larger slope than the green line (2000 mm/min) over different pressures.The data for the explored parameter ranges fail to reject the hypothesis that there is no significant interaction between the varied parameters. This indicates that as speed increases and pressure decreases, the line feature width decreases. On VeroWhite+, the smallest achievable trace was 879 µm and 720 µm on TangoBlack+.

The relationships that exist are:• • Features on TangoBlack+ are narrower than those on VeroWhite+Increased pressure yields wider features Increased toolhead speed yields narrower feature widthIf a consistent bead width is critical across the different substrates, the tool path could specify a change in toolhead speed to maintain consistency or adjust the dispense pressure. For example, it is shown that dispensing on TangoBlack+ at 37.92 kPa and 2000 mm/min, the feature width is comparable to dispensing on VeroWhite+ at 31.03 kPa and 3000 mm/min. The bead width differs by only 24 µm on average. Additionally, this feature width data can be used to assist with designing channels into the PolyJet parts for full encapsulation of the deposited conductive trace.For all remaining experiments in this work, a dispense pressure of 31.03 kPa and a toolhead speed of 3000 are used since these settings yield the smallest feature sizes.

While the bead height is not characterized in this experiment, it is measured in Section 3. The rationale is that these width measurements are taken before the ink is fully dry on the PolyJet substrate. Optical measurement indicated that the drying process does not affect the width, but the height changes considerably as the solvent evaporates. The next section discusses how different drying conditions affect the ink geometry. Once the drying process is better understood, height data is measured and discussed.4 EFFECTS OF DRYING CONDITIONS ON DEPOSITED TRACE HEIGHT4.1 ContextThe drying conditions of conductive inks are understood to have a positive effect on the conductivity and adhesion of conductive materials [15]. Given the motivation of creating functional circuitry, high conductivity and strong adhesion is desirable. Commercially available conductive inks are prescribed a drying regimen of specific temperature and duration for ensuring quality conductivity and adhesion.

Commercially available conductive inks are prescribed a drying regimen of specific temperature and duration for ensuring quality conductivity and adhesion. It is recommended by the supplier to cure the DuPont 5021 ink at 120°C for at least 5 minutes [15]. However, these temperatures are in excess of the PolyJet material’s glass transition temperatures (54°C) and, therefore, not compatible.This experiment aims to explore how varying dry times and temperatures (compatible with PolyJet material) affect the final conductivity and adhesion of silver-loaded conductive ink on PolyJet substrates.

More simply, how does drying time and temperature affect the geometry of conductive inks on PolyJet substrates? In addition, the height of dried lines is analyzed to determine if the narrower lines observed on TangoBlack+ versus VeroWhite+ substrates result in taller features.Initially, it is hypothesized that drying the ink at temperatures near the published glass transition temperatures of the PolyJet materials will not significantly impact adhesion and profile geometry of the conductive traces. It is observed that as the deposited beads dry, the cross- section settles and shrinks in the Z-direction as the solvent evaporates. Rapid drying could7potentially inhibit this settling effect and cause different cross-sectional profiles than specimens dried over a longer period of time at room temperature.

Rapid drying could7potentially inhibit this settling effect and cause different cross-sectional profiles than specimens dried over a longer period of time at room temperature. If the geometry is not significantly affected, the heated drying regimen can be used as higher temperatures help the ink cure faster.4.2 MethodThe experiment outlined in this section is designed to investigate the effects of drying time and temperature on the geometry of the conductive traces on the PolyJet substrate. In order to understand the potential effects on adhesion on the cross-sectional geometry, contact angle and bead height are measured. Two sets of samples were evaluated: (i) samples that are allowed to dry in air for 24 hours and (ii) samples that are dried at 55 °C for 30 minutes (which is near the material’s heat deflection temperature).

As discussed in Section 2, the heated samples are exposed to heat mid-print on the PolyJet build tray using a heat gun.3-dimensional profilometry (A HIROX Digital Video Microscope) is used to determine if there is any significant change in the cross-sectional profile between the two substrates’ sample sets. The microscope uses a composite of multiple images with different focal planes to construct a 3D profile as shown in Figure 2. From the 3D profile, cross-sectional profile and height data is made available.Figure 2: Demonstration of HIROX multi-focus 3D synthesis process (used under fair use, 2013[16])For this experiment, conductive traces were deposited on the surface of the PolyJet VeroWhite+ and TangoBlack+ substrates using the dispense parameters determined in Section 3.2: the dispense pressure was set to 31.03 kPa (4.5 psi) and the toolhead speed was 3000 mm/min using a 406 µm diameter dispensing tip.

Once the traces were deposited, small sections were viewed under the microscope. A demonstration of where the microscope is focused is shown in Figure 3.8(a)(b)Figure 3: Demonstration of sample profilometry locationIn Figure 4, topographical data of only half the bead is visible because at the magnification necessary to gather 3-dimensional data from different focal planes the field of view is limited. However, it is assumed that the bead’s profile is symmetric. Using the HIROX multifocal synthesis, a 3D model is constructed as shown in Figure 4.(a)(b)Figure 4: 3-dimensional model using HIROX multi focal synthesis with (a) actual image data overlaid and (b) a colored surface for indicating relative height.4.3 ResultsWith five deposited traces for each material (TangoBlack+ and VeroWhite+), and sampling 5 maximum heights from each trace, a total of 50 height measurements are recorded.

The maximum height of each deposited trace was analyzed for the different samples that underwent room temperature drying and heated drying on both TangoBlack+ and VeroWhite+ substrates. The recorded data is tabulated in Table 4. The data shows only a small difference in the observed average heights between those dried at room temperature and those accelerated with heat. On the samples without any sort of heat treatment, the heights averaged to be 27 and 36 µm on VeroWhite+ and TangoBlack+, respectively. With heat treatment, there was very little change. The average heights were 28 µm on VeroWhite+ and 38 µm on TangoBlack+.Using a simple t-test, assuming a two-tailed distribution and an alpha of 0.05, a p-value of 0.583 was calculated; we therefore cannot conclude a significant difference in geometry between the drying methods on either TangoBlack+ or VeroWhite+.

The accelerated drying conditions do not significantly affect the profile geometry of the conductive materials to the PolyJet substrate.9Table 4: Tabulated height data for samples (i) heated at 55 °C for 30 minutes and (ii) air- dried for 24 hoursSampleHeat	
 TreatmentHeight	
 [µm]VeroWhite+	
 TangoBlack+Average	
 [µm]Standard	
 Dev.	
 [µm]1	
 2Y	
 Y26.06	
 27.6638.55	
 37.48VeroWhite+27.5130.9193	
 4Y	
 Y28.39	
 26.9437.86	
 37.32TangoBlack+5Y28.5137.3637.7140.4596N26.1735.45VeroWhite+7N31.4136.1726.7582.5488	
 9	
 10N	
 N	
 N24.06	
 27.14	
 25.0136.22	
 35.93	
 37.14TangoBlack+36.1820.5514.4 DiscussionThis experiment shows that the heating process, which helps the ink dry faster, does not significantly impact the final geometry of the deposited features.

There is no information from the supplier that indicates temperatures below the recommended curing temperature will have any positive impact on the final conductivity [3]. The heat applied is to reduce the drying time before printing can be resumed. The height of the traces after heating is also of interest because it can influence the decision to design channels for embedding the material into the printed part. Traces on VeroWhite+ are shorter than 32 µm, which is the layer height of the PolyJet process. Traces on TangoBlack+ are only 5 µm taller than the PolyJet layer height on average. Because of this, channels should be designed to be 32 µm deep as the PolyJet process would not be able to accommodate shallower features.Reflecting on results shown in Section 3 regarding profile geometry, we also notice a difference in heights between the VeroWhite+ and TangoBlack+ features. Recalling the results from Experiment 1, we know that the features on TangoBlack+ are narrower.

Recalling the results from Experiment 1, we know that the features on TangoBlack+ are narrower. Therefore, from Equation 1, the height should also be taller. This relationship is confirmed by these results. Conductive features on TangoBlack+ are taller and narrower than those on VeroWhite+. This effect is possibly related to adhesion or surface interaction of the conductive materials to the PolyJet substrates and is explored in Section 5.5 EFFECT OF POLYJET SUBSTRATE ON CONDUCTIVE INK ADHESIONAs mentioned in Section 1, the embedding process prevents the servicing of integrated conductive traces after manufacture. An internal connection failure or physical break in the conductive trace could render the component nonfunctional. In order to ensure reliable connections, it is important to understand the adhesion characteristics between the conductive materials and the PolyJet substrates.

In order to ensure reliable connections, it is important to understand the adhesion characteristics between the conductive materials and the PolyJet substrates. When not bonded to a surface, the thin conductive traces10are prone to breaks from movement or vibrations. However, when the conductive material is able to bond to a surface, the traces are constrained and less susceptible to breaks. Good bonding between the conductive material and PolyJet substrate creates stable traces and thus reliable connections. Poor bonding, results in suspended traces that are more susceptible to cracking and failure.Material wetting is a fair indicator of inter-material adhesion [17]. Material wetting is defined as how well a liquid phase material contacts a solid surface on a molecular level and is a result of interaction of cohesive and adhesive forces.

Material wetting is defined as how well a liquid phase material contacts a solid surface on a molecular level and is a result of interaction of cohesive and adhesive forces. Young’s equation (Equation 2) effectively models the scenario of a fluid droplet on a solid surface by relating solid-liquid free energy (𝛾!"), solid- surface free energy (𝛾!"), and liquid-surface free energy (𝛾!") to the droplet’s contact angle.Materials with contact angles below 90° are considered hydrophilic and exhibit a strong solid-liquid interaction. This interaction provides ample contact between the liquid and solid for adhesion.

This interaction provides ample contact between the liquid and solid for adhesion. The experiments designed to measure the contact angle to determine if the conductive inks adhere well to the PolyJet substrates.𝛾!" = 𝛾!" + 𝛾!" cos 𝜃𝜃 = Contact Angle𝛾!" = liquid-surface free energy𝛾!" = solid-surface free energy𝛾!" = solid-liquid free energyFigure 5: Diagram of Young’s equation variables5.1 MethodSample pieces of VeroWhite+ and TangoBlack+ substrates were printed and droplets of the conductive material were applied. The droplets were semi-manually applied to the substrate using a discrete 0.25-second extrusion dosage from the Nordson EFD pneumatic dispenser, as shown in Figure 6. In order to measure the contact angle, a digital microscope was positioned parallel to the adhesion surface and the sample was backlit in order to create a silhouetted droplet as shown in Figure 7F.

In order to measure the contact angle, a digital microscope was positioned parallel to the adhesion surface and the sample was backlit in order to create a silhouetted droplet as shown in Figure 7F. Using the images captured by the microscope, the contact angle was digitally measured via ImageJ measuring software. A sample image is shown in Figure 8. Finally, transparent pressure-sensitive tape is applied and removed to the dried sample to see if any material is removed to test adhesion strength.Figure 6: Droplets produced by 0.25 second dispensing pulses on VeroWhitePlus substrate11(2)Figure 7: Contact angle measurement setup schematicFigure 8: Example of ImageJ digital measuring process5.2 ResultsThe resultant contact angle measurements are presented in Table 5. Droplets on the VeroWhite+ substrates had an average contact angle of 31.6 degrees with a standard deviation of 2.7 degrees.

Droplets on the VeroWhite+ substrates had an average contact angle of 31.6 degrees with a standard deviation of 2.7 degrees. The TangoBlack+ samples had an average of 37.7 degrees with a standard deviation of 3.2 degrees. Both of these averages are well within the hydrophilic range and demonstrate good wettability and adhesion. Moreover, no material was removed from either substrate using the scotch tape test. When adhered and removed, no material was visibly resent on the surface of the tape.As hypothesized from observations in Experiment 1, the conductive material on the TangoBlack+ surfaces averaged a larger contact angle, which would imply that the lines are thinner for a constant cross-sectional area.

An example of the VeroWhite+ and TangoBlack+ profiles that demonstrate these differences are shown in Figure 9.(a) VeroWhite+(b) TangoBlack+Figure 9: Comparison of contact angles on (a) VeroWhite and (b) TangoBlack substrates12Table 5: Contact angle measurementsContact	
 Angle	
 [°]VeroWhiteTangoBlack131.933.74233.236.99334.740.02435.1539.47532.8236.3631.1235.4731.5944.5830.5439.75930.2236.181025.3335.04Average31.6637.74Std.	
 Dev.2.773.25.3 DiscussionThe contact angle results provide further insight into the results from Section 3 that demonstrated that features on TangoBlack+ are thinner than on VeroWhite+. Given that both average contact angles are well within the hydrophilic range, good adhesion is expected on both substrates, although adhesion to VeroWhite+ is relatively better (as per the smaller contact angle).

Additionally, ink on both substrates passed the “scotch tape test.”While using the HIROX digital video microscope during experiments conducted in Section 4, it was observed that the PolyJet material surfaces (both VeroWhite+ and TangoBlack+) had many small imperfections or divots on the surface, as shown in Figure 10. It is unclear what causes the small holes; it may be attributed to inkjet head misfires, imperfections in the roller, or adhesion to the roller. However, these holes may also play an important role in the conductive materials’ adhesion to the PolyJet surface on the macro scale by allowing the material to penetrate and gain more surface area for adhesion.13Figure 10: Microscope image of surface imperfections on TangoBlack+ substrates.6 CLOSUREWith a goal of directly fabricating mechatronic products with integrated actuation and sensing, the authors are exploring the integration of extrusion-based DW with PolyJet material jetting AM.

As discussed in Section 2, this goal has been achieved via the development of an extrusion-based DW system that can be placed directly into the PolyJet build tray to enable the deposition of conductive materials (DuPont 5021 silver-loaded conductive ink) directly onto in- situ 3D printed parts.The goal of this paper is to gain an understanding of how process parameters affect material compatibility of the DW and PolyJet systems. Specifically, experiments were conducted to determine (i) how DW extrusion parameters (tool tip diameter, toolhead speed, and extrusion pressure) affect the deposited bead width on two PolyJet materials, (ii) how post- processing drying parameters affect the deposited bead height, and (iii) how the adhesion characteristics of the deposited material differ on two different PolyJet substrates.

Regarding (i), it was found that increased pressure yields wider traces, increased toolhead speeds yield narrower traces, and deposited traces on TangoBlack+ are narrower than those by VeroWhite+ (Section 3). Regarding (ii), it was observed that the use of a heat gun does improve material drying times and does not affect the geometry of the trace profile (Section 4). Regarding (iii), it was discovered that the conductive material deposited onto the TangoBlack+ material had a larger contact angle than that deposited on the VeroWhite+ material, which explains the thinner traces. Furthermore, contact angles of deposits on both substrates were well within the hydrophilic range, which suggests good adhesion (Section 5). These results provide detailed characterization of the profiles of the deposited conductive ink, which can to design channels into the PolyJet part for full encapsulation of a deposited conductive trace.

Future work will focus on exploring (i) how conductivity changes as a function of temperature in the drying process and, (ii) how conductivity is affected by full encapsulation.147 ACKNOWLEDGEMENTSThe authors acknowledge support from the Manufacturing and Industrial Technologies Division of the Air Force Research Laboratory Materials and Manufacturing Directorate, which directly supported this work.8 REFERENCES[1]J. A. Palmer, P. Yang, D. W. Davis, B. . Chavez, P. L. Gallegos, R. B. Wicker, and F. R. Medina, “Rapid Prototyping of High Density Circuitry,” no. 313.[2]F. Medina and A. Lopes, “Hybrid Manufacturing: Integrating Direct-Write and Stereolithography,” Proc. …, pp. 39–49, 2005.[3]A. J. Lopes and U. Consolidation, “Integration of Direct-Write ( DW ) and Ultrasonic Consolidation ( UC ) Technologies to Create Advanced Structures with Embedded Electrical Circuitry .,” pp. 60–69, 2006.[4]E. Malone, M. Berry, and H.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCHARACTERIZATION OF LASER DIRECT DEPOSITED MAGNESIUM ALUMINATE SPINEL CERAMICSJohn M. Pappas, Xiangyang Dong*Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409.AbstractAn additive manufacturing (AM) approach, via laser direct deposition, is investigated in printing transparent magnesium aluminate spinel (MAS) ceramics. Using AM, traditionally difficult or expensive to manufacture shapes, such as optical lenses, can be rapidly manufactured to near net shape, reducing time consuming and expensive post processing requirements. The transparency of MAS ceramics is highly dependent on the microstructure, with porosity and microcracking having the largest effect on the transparency of fabricated parts.

The transparency of MAS ceramics is highly dependent on the microstructure, with porosity and microcracking having the largest effect on the transparency of fabricated parts. With high localized heat inherent in the laser deposition process, the microstructure of ceramic parts can be controlled by adjusting processing parameters. In this study, thin wall MAS structures were fabricated by varying processing parameters. Processing parameters including laser scan speed and laser power had a large influence on the part quality. To fabricate transparent magnesium aluminate spinel ceramics with high mechanical properties, the effects of processing parameters on part porosity, density, and microstructure were studied. Dense MAS parts were successfully fabricated through the laser direct deposition process. Low scan speed and high laser power showed the most promising results in fabricating MAS parts of low porosity.

Low scan speed and high laser power showed the most promising results in fabricating MAS parts of low porosity. Using a low powder flow rates of 0.58 g/min, a relative density of nearly 98% was achieved. Directional cooling through the substrate and from the powder conveying gas led to columnar grain growth at a tilt angle from the build direction. The primary defects of fabricated MAS ceramics were found to be residual porosity and microcracking, which negatively affected part transparency and mechanical properties. Typical microcracking patterns included transverse and longitudinal cracking, with longitudinal cracks being more prevalent due to the existence of columnar grains and intergranular fracture mode.

Typical microcracking patterns included transverse and longitudinal cracking, with longitudinal cracks being more prevalent due to the existence of columnar grains and intergranular fracture mode. A preliminary study demonstrated that a certain degree of transparency was achieved in additively manufactured MAS ceramic parts via laser direct deposition.Keywords: Transparent ceramics, Magnesium aluminate spinel, Laser direct deposition, Additive manufacturingIntroductionTransparent ceramics have many useful applications including transparent armor, lenses, and other optical elements. Among them, magnesium aluminate spinel (MAS) ceramics are of particular interest due to high in-line transmittance [1] and the highest hardness compared to other transparent materials [2], along with other desirable properties including cubic crystalline structure, which is optically isotropic and has high transmission values through a wide variety of grain sizes [3].

Non-cubic polycrystalline ceramics, such as α-alumina, are much more difficult to manufacture with high transparency due to birefringent nature, where light is scattered by grainCorresponding Author: dongxi@mst.edu, 1-573-341-4373 (ph), 1-573-341-4607 (fax)907boundaries [4], and non-cubic polycrystalline ceramics must have sub-micrometer grain size to be transparent [5]. In particular, considering the challenges still encountered in controlling grain size and orientation via additive manufacturing (AM) [6], cubic polycrystalline ceramics, including MAS, work as a better alternative in fabrication of transparent parts over non-cubic polycrystalline ceramics for the laser direct deposition process described in this paper.Previous studies showed that a major challenge of laser deposition-based AM was the difficulty in fabricating fully dense parts due to residual porosity.

Porosity negatively affects the optical properties of transparent ceramics through radiation scattering mechanisms [5-6]. Residual porosity is the main cause of reduced transparency of fabricated parts, especially in the visible and ultraviolet range [9]. Depending on the number of pores and void size, fabricated parts can either exhibit opaque or translucent characteristics instead of true transparency. If pore size is close to the optical wavelength, transmittance through the ceramic is at a minimum, with small pores showing much less optical scattering than larger pores [10]. To be transparent, materials must have a residual porosity of <0.1% with pore sizes less than 100nm [1]. To have very high-glass like transmittance, residual porosity must be further reduced to <0.01% [11]. Porosity formation is highly process dependent; by adjusting processing parameters, the prevalence and size of pores can be affected to a large extent. Mahamood et al.

Porosity formation is highly process dependent; by adjusting processing parameters, the prevalence and size of pores can be affected to a large extent. Mahamood et al. [12] found that laser power and scan speed played significant roles in residual porosity. It was shown that with increased laser power, and decreased scan speed, porosity was reduced. Porosity-free Ti6Al4V parts could be obtained at a laser power of 800 W and a scan speed of 5 mm/min fabricated by laser metal deposition. Ning et al. [13] found that using an ultrasonic vibration table, porosity was significantly reduced for Inconel 718 parts fabricated with the laser engineered net shaping process.

Ning et al. [13] found that using an ultrasonic vibration table, porosity was significantly reduced for Inconel 718 parts fabricated with the laser engineered net shaping process. This was attributed to the acoustic streaming and cavitation effect caused by ultrasonic vibrations.In addition to porosity, microstructural characterization and control of resultant microcracking, grain size, and grain distribution are very important due to the potential impacts they have on transparency and mechanical properties of fabricated parts. Cracks severely weaken the mechanical properties and negatively affect the transparency of parts. Therefore, research work has been done to minimize or eliminate process-induced cracks as much as possible. Niu et al. [14] found that crack formation and prevalence was significantly affected by layer thickness for pure alumina structures fabricated with laser engineered net shaping.

Niu et al. [14] found that crack formation and prevalence was significantly affected by layer thickness for pure alumina structures fabricated with laser engineered net shaping. In this study, it was found that crack prevalence significantly decreased with increasing layer thickness, and at layer thickness of 0.8 mm and above, nearly crack-free specimens were produced. In a related study, Niu et al. [15] found that increasing laser scan speed decreased the number of cracks in pure alumina ceramics produced by laser engineered net shaping, and a scan speeds greater than 700 mm/min produced crack-free structures. This result was attributed to increased scan speed decreasing thermal stress and increasing fracture strength. Inclusion of ultrasonic-vibration during laser engineered net shaping of alumina-zirconia parts was studied by Hu et al. [16].

Inclusion of ultrasonic-vibration during laser engineered net shaping of alumina-zirconia parts was studied by Hu et al. [16]. It was found that crack initiation and propagation were suppressed due to the addition of ultrasonic vibration, which was primarily attributed to grain refinement and reduced thermal stress. Additional benefits of using ultrasonic vibration during fabrication were increased mechanical properties including microhardness, wear resistance, and compressive strength. Balla et al. [17] found that alumina ceramics fabricated with laser engineered net shaping showed anisotropic mechanical properties due to columnar grain solidification along the build direction, with marginal increases in strength after additional high- heat treatment. Niu et al [18] found that columnar grains in pure alumina ceramics produced by908laser engineered net shaping are typically larger in ceramics produced by sintering methods, and metals prepared with the same method.

It was proposed that a primary reason for this was due to the low thermal conductivity of ceramics causing heat accumulation during fabrication.With a focus on transparent ceramics, the goal of this work is to study the effects of processing parameters and powder precursors on part residual porosity and microstructure as well as investigate the viability of laser direct deposition for MAS. Due to the critical role that porosity plays in transparency of these ceramics, the effect of processing conditions on resultant porosity was a key focus in this paper. While there has been considerable work on AM of non-transparent ceramics via selective laser melting (SLM) [19]–[22] and laser engineered net shaping [16], [18], [23], [24], there is a lack of work on transparent ceramics, including MAS. This could be attributed to difficulties in additively manufacturing MAS ceramics due to process related porosity and cracking.

This could be attributed to difficulties in additively manufacturing MAS ceramics due to process related porosity and cracking. Therefore, it is necessary to characterize the microstructure of direct laser deposited MAS ceramics and determine which processing conditions mostly affect residual porosity on the fabricated parts.Experimental Procedure2.1 Powder PreparationThe materials in this study included MgO powders purchased from Sigma-Aldrich (Part No. 342793) with a purity greater than 99% and average particle sizes less than 44 μm. The MgO powders were used with Al2O3 powder from Almatis (A3500 UG) with a purity of greater than 99.5% and D-50 particle size of about 2.2 μm. An excellent flowability was observed for the selected powders during the fabrication processes in this study. Following the stoichiometric ratio of MAS, the powder was first mixed in a ball mill for two hours using alumina grinding media and acetone solvent.

Following the stoichiometric ratio of MAS, the powder was first mixed in a ball mill for two hours using alumina grinding media and acetone solvent. Acetone solvent was removed with a Buchi R124 rotary evaporator. Following the rotary evaporation, the powder was calcined with forced air flow for eight hours in a Lindberg furnace at 600 °C to burn out any organic contamination, especially residual acetone and dehydrate the MgO powder.It is worth noting that in this study, the MAS ceramics were directly synthesized from the Al2O3-MgO powder mixtures via melt-grown methods during laser direct deposition. This method of synthesis was chosen due to the difficulties in purchasing commercially available MAS powders that showed consistent and repeatable flowability required in the powder feeding process for laser deposition in this study. Therefore, alumina and magnesia powders that showed good flowability were first chosen and mixed as described previously.

Therefore, alumina and magnesia powders that showed good flowability were first chosen and mixed as described previously. Due to low availability of spinel powders, options are very limited with respect to particle shape and size, both of which significantly influence how powders will flow in the powder feeding processes. Preliminary studies were performed using the MAS powder from Sigma-Aldrich with an average particle size of < 50 nm. Testing with this powder proved the difficulty of maintaining a consistent powder flow, which resulted in very poor deposition quality, as depicted in Figure 1. In addition to the improved flowability needed in powder feeding processes, the synthesis method used in this study showed greater potentials in reducing the manufacturing costs, considering the availability of magnesia and alumina powders separately.

In contrast, commercially available spinel powders are909commonly more expensive and difficult to obtain at the specific stoichiometric composition for MAS.Balling~Inconsistent DepositionIFigure 1: Image showing poor deposition quality of commercially available MAS powders due to inconsistent powder flow, which illustrates the main benefit of the laser-melting method of spinel synthesis from magnesia and alumina powder mixtures.It has also been observed that a potential issue with this synthesis process was related to MgO, which is a hygroscopic substance and readily reacts with water vapor in the air to form magnesium hydroxide (Mg(OH)2). If the powder was exposed to moisture before processing, the flowability of the powder was seen to be severely limited, making it difficult to accurately control the powder flow rate. Hydrated powder may introduce additional gasses into laser processing, thus increasing residual porosity in fabricated parts.

Hydrated powder may introduce additional gasses into laser processing, thus increasing residual porosity in fabricated parts. Decomposition of Mg(OH)2 occurs at temperatures greater than 450 °C [27], [28], thus requiring calcination immediately prior laser deposition. This necessitates the need of careful preparation of Al2O3-MgO powder mixtures with a burn-out procedure added in this study.2.2 Experimental Setup and CharacterizationCoorsTek AD-998 alumina substrates with 99.8% purity were used in this study for thermal expansion compatibility with deposited MAS. The laser direct deposition setup used in this study is shown in Figure 2 and includes a Convergent Energy Arrow Ultimate CO2 laser with maximum continuous wave (CW) output of 1.7 kW at 10.6 μm wavelength. The alumina substrate was mounted onto a 3-axis motorized table for positioning and movement.

The alumina substrate was mounted onto a 3-axis motorized table for positioning and movement. Powder delivery was controlled with a Powder Motion Labs X2W powder feeder and controller exhibiting a feed rate repeatability of less than 1%. Argon inert gas was used to feed the prepared powder mixture at a flow rate of 3 L/min.

Argon inert gas was used to feed the prepared powder mixture at a flow rate of 3 L/min. A 1.6 mm inner diameter alumina tube was used to guide the powder into the melt pool parallel to the scan direction and at an angle from the substrate.910Laser System ----1 I I I I I I I I I IIPo;;:r ~iv::; 1- -I 1 Beam Delive1ySystemGas MixerCO2 Laser GeneratorL _ system 1 IPowder FeederLaser Power Supply---, I Laser Focusing IPowderFeed TuberArgonPowder Delive1yLens & Z-axisPositioning ILaser ChillerNozzle - - -Deposition r I I I Powder I L~- ..lChamberConvey ingX-Y Positioning TableFigure 2: A schematic diagram of the experimental setup for laser direct depositionThe processing parameters used for this study are summarized in Table 1. Laser spot size was maintained at a constant diameter of 2.5 mm in this study. Parametric studies were performed by varying laser power and scan speed to study their effects on the printed ceramic parts.

Parametric studies were performed by varying laser power and scan speed to study their effects on the printed ceramic parts. Thin- wall structures were built using a reciprocal scanning motion. After each track, the laser head would increment in the z-direction by Δz, which was equivalent to the layer thickness.Table 1: Process parameters used in this studyLaser Power (W)275, 485Laser Spot Size (mm)2.5Scan Speed (mm/minute)1000, 2000, 3000, 4000, 5000The printed thin wall samples were then sectioned and polished for microstructural characterization. The images of the polished samples were captured using a Hirox KH-8700 digital microscope. Due to limitations in the field of view of the optical microscope, image stitching was performed in Fiji software using an image stitching plugin [29]. Due to the variations in processing conditions for the top and bottom layers, the middle sectioned were analyzed to obtain the typical microstructural characteristics.

Due to the variations in processing conditions for the top and bottom layers, the middle sectioned were analyzed to obtain the typical microstructural characteristics. ImageJ software was used to analyze the porosity area of the fabricated MAS samples, and Archimedes’ method was used for the density analysis of as- fabricated thin wall structures. To reveal the typical grain size and distribution, the polished structures were thermally etched in accordance with the guidelines in the ASM Handbook Volume 9 [30], at 1400 °C for 10 minutes.911Results and Discussion3.1 Microstructure Characteristics of MAS SamplesFigure 3 shows the typical microstructure characteristics observed in the printed MAS samples after thermal etching. Columnar grains were primarily observed for MAS samples fabricated by laser direct deposition. Figure 3B shows the cross-sectional view of the whole sample and highlights areas taken for closer microstructural examination.

Figure 3B shows the cross-sectional view of the whole sample and highlights areas taken for closer microstructural examination. Figure 3A shows the microstructure of a region deposited near the substrate. Initially deposited layers exhibited columnar grains grown at a tilt angle of approximately 20 degrees from the build direction. This was mainly due to the solidification principal of columnar grain growth occurring in the orientation of thermal gradients, opposite to the direction of heat flow, where rapid cooling rates are present [17]. The 20-degree grain orientation with respect to the build direction was attributed to the primary cooling mechanism being conducted through the substrate, with lesser convective cooling rates by the argon powder conveying gas.

Hence, the largest thermal gradient for initially deposited layers was along the build direction, resulting in the observed columnar grains.Figure 3: Typical microstructure of thin wall sample fabricated at 275 W laser power, 2000 mm/min scan speed, and 2.1 g/min powder flow rate by laser direct deposition showing: A. Columnar grains grown at a small angle from the build direction for the layers deposited near the substrate; B. Cross-sectional view of the entire sample; C. Columnar grains grown at a larger angle from the build direction for the layers deposited further from the substrateFigure 4 depicts the side view of thin wall structures and illustrates the effects of the conveying gas plume around the thin wall structure on the side adjacent to the powder feed nozzle. The powder feed nozzle was aimed at the melt pool perpendicular to the scan direction, and at an angle from the build direction in this study.

The powder feed nozzle was aimed at the melt pool perpendicular to the scan direction, and at an angle from the build direction in this study. This causes a continuous flow of cool argon gas around the outer surface of the deposited structure on the right side, thus producing relatively high cooling rates, and hence thermal gradients, on one side. The opposite side was mostly blocked from the912argon flow by the structure itself and therefore experienced much lower cooling rates, which are sufficient to significantly affect grain growth.It is also interesting to note that the layers deposited further from the substrate had an increased tilt angle of approximately 40 degrees from the build direction as shown in Figure 3C. Since the MAS ceramics were a good type of insulator materials, heat flow to the substrate reduced considerably at an elevated distance from the substrate.

Since the MAS ceramics were a good type of insulator materials, heat flow to the substrate reduced considerably at an elevated distance from the substrate. Therefore, the percentage of total heat loss to the side of the structure, due to convection from the powder conveying gas, increased. With this increase in transverse heat flow, grain orientation shift away from the build direction increasingly with an increased distance from the substrate.Powder Feed NozzlePowder Conveying Gas PlumeThin Wall Strncture Side ViewConvective Cooling on the Side of Deposited LayersSubstrateFigure 4: Schematics depicting convective cooling due to the powder conveying gas plume on one side of thin wall structures during fabrication3.2 Cracking Characterization of MAS SamplesThe primary microcrack propagation directions for MAS thin wall structures are in the build direction (longitudinal cracks) and along the scan direction (transverse cracks).

Figure 5 shows a backlit sample to highlight typical cracks observed. Very high cooling rates due to laser melting and solidification processes, of up to 105 K/s [17], create large thermal gradients and thermal stresses in the deposited materials. The relatively high hardness of the deposited MAS ceramic material makes it very susceptible to brittle fracture, which occurs when the tensile stresses are higher than the mechanical strength [31].913It was observed that longitudinal cracks are more prevalent, and typically extended further than transverse cracks. It could be attributed to the columnar grains extended primarily along the longitudinal (build) direction. As higher fracture energy was often required for transgranular fracture, intergranular cracking would be more common in the brittle materials and thus primarily propagated along the grain boundaries of vertically elongated columnar grains [14].

On the other hand, the cooling and solidification processes of recently deposited layers led volume contraction, which was constrained by the previously solidified but cooler layers [15], thus yielding tensile stresses in the newly deposited layer. The tensile stresses initiated microcracks, which preferentially propagated along the grain boundaries as discussed above.Figure 5 Side view of a MAS sample showing significant cracking in the longitudinal direction with fewer transverse cracksTransverse cracks (along the scan direction) were observed mostly near the sides of thin wall structures, extending up to 5 mm along the scan direction. Transverse cracks initiated at defects typically located near the edges of the deposited thin wall structures. Immediately after deposition, the part of the melt pool at the interaction zone with the previous layer solidified first due to the high conduction rates through the lower layers.

Immediately after deposition, the part of the melt pool at the interaction zone with the previous layer solidified first due to the high conduction rates through the lower layers. Volume contraction, during solidification, of the upper portion of the melt pool creates a bending moment on these layers resulting in tensile forces parallel to the build direction at the sides of thin wall structures [15]. Cracks would be formed if the tensile stresses were higher than the strength of the newly deposited914material. However, transverse crack length was typically limited due to crack bifurcation into the preferential fracture direction, i.e., along the grain boundary of columnar structures as discussed above. Further studies will be performed to investigate the effects of processing parameters on crack prevalence within the deposited MAS ceramics.Transverse cracks were seen to be susceptible to crack bifurcation, but also crack deflection guided by residual pores.

. : ' ; .. :: ~~ ;~---~. :· / .'~:; :': ::} . ~.;:I·.. ..C\ ~, ... ~.. ' ~ "'·... . . .. ·,.,. '. •. . a. \ ... -..· • . :, • ,: . ., ' . -.' '• . I1f .._-j- • ... \'•500 fllllFigure 6: Cross-sectional view of MAS sample showing a transverse crack with crack deflection due to residual pores and crack bifurcationLongitudinal cracks commonly occurred near the substrate of the printed samples as shown in Figure 7. This was mainly due to the high temperature gradient near the substrate-deposition interaction zone. The solid substrate was at a lower temperature than the deposited material, and therefore contracted considerably less during cooling processes. Thus, the substrate would undergo compression stresses while the deposition layers was under tensile stress. As a result, the tensile stresses exceeded the strength of the deposited material causing crack initiation and propagation to occur at the substrate-deposition interaction zone.

As a result, the tensile stresses exceeded the strength of the deposited material causing crack initiation and propagation to occur at the substrate-deposition interaction zone. The longitudinal cracks present at the substrate-deposition interface typically only extended approximately 1 mm in the build direction since the temperature differential, and therefore tensile stress, decreased with an increasing distance from the substrate.Another consequence of the high temperature differential at the substrate-deposition interaction zone was that the deposited material experienced much more volume contraction than the substrate during cooling. The volume contraction causes a bending moment to run through the sample, and as a result, high tensile forces were present along the sides of the deposited material915near the substrate-deposition interaction zone.

Occasionally, single and long cracks were observed in the deposited samples as shown in Figure 7, which were deflected by the residual pores. Tensile forces generated were also typically large enough to cause brittle failure between deposited layers and the substrate and frequently led to the delamination of the thin wall samples from the substrate. Due to the bending moment, warping of the deposited material was typically seen, especially in long thin wall structures (length >25 mm) in the preliminary tests. To minimize or even eliminate the longitudinal cracks near the substrate, as well as the delamination of the deposited part, it is suggested to preheat the substrate prior to the deposition process. Sufficient substrate preheating would significantly reduce thermal gradients in this region and thus reduce tensile stresses and bending moment in the samples.. ~- \ .·. . -· . ··-: ,-_ -• ~ . .... ... .. ' • .. . ..,. . \ .,.._' : ·.·,,. ... . -~ . ~ .. ~- . ... :· . ..

~- \ .·. . -· . ··-: ,-_ -• ~ . .... ... .. ' • .. . ..,. . \ .,.._' : ·.·,,. ... . -~ . ~ .. ~- . ... :· . .. - ~~ .... ··' ... .. ..Figure 7: The optical micrograph of the polished sample side view showing severe longitudinal cracking near substrate-deposition interface due to contraction induced tensile stresses in the deposited MAS structure3.3 Porosity Characterization of MAS Samples3.3.1 Effects of Laser Scan Speed on Residual PorosityThe effect of scan speed on the porosity of fabricated parts was investigated and Figure 8 shows the primary results. The general trend for scan speed with respect to porosity was that as scan speed increased, there was an increase in porosity as shown in Figure 8A. The minimum porosity was 9.8% and the maximum was 14.5% at scan speeds of 1000 mm/min and 5000 mm/min, respectively.

The minimum porosity was 9.8% and the maximum was 14.5% at scan speeds of 1000 mm/min and 5000 mm/min, respectively. The results for the density of the entire structure can be seen in Figure 8B.916The highest average density was 96.6%, while the lowest was 92.5% at scan speeds of 4000 mm/min and 1000 mm/min, respectively. The density analysis of the entire part showed a clear trend, where increased scan speed resulted in increased relative density.A.B .20100,-._ ~ c:,:i 15 V$ .clO ·;;; 0 ;.. 0 p.. 5C c:s:l -~ .....V ;.. 0 V95~ 90 ~ ..__,, _g .::: Q) Cl85r/l080C.020004000Scan Speed (mm/min)6000D .02000 Scan Speed (mm)40006000100,-._ ~ ..__,, _g .::: V Cl V .:;! ..... c:s:l vA95 ~ r/l90A~85800200040006000Scan Speed (nun/min)Figure 8 Results from scan study on A. Porosity area and B. Sample density, also showing C. Thin wall sample ground to cuboid shape, and D.

Porosity area and B. Sample density, also showing C. Thin wall sample ground to cuboid shape, and D. Density measurements of the post processed samples.It should be noted that this trend was not correlated well with the porosity results for this study. It was expected that an increase in density would be related to a reduction in porosity. It is important to note that open porosity may have significantly affected measured density of samples showing higher porosity. As the frequency of pores increased, there was a greater chance that pores would connect. Connected pores that were open to the surface would fill with distilled water used during density testing, resulting in artificially high measured densities. Samples with lower porosity levels, and hence less open porosity, would have a smaller interior volume filled with distilled water, resulting in lower but more accurate measured densities.

Samples with lower porosity levels, and hence less open porosity, would have a smaller interior volume filled with distilled water, resulting in lower but more accurate measured densities. Increased open porosity could help to explain why density increased while porosity increased. The problem with measuring the density for the entire sample was that it may not be representative of the usable central portion. Since the area used for porosity analysis only represented a portion of the entire sample, it is possible that the porosity results could vary considerably depending on where the cross-section is917taken. When fabricating parts with this method, post processing would be necessary to remove the unstable regions which typically show atypical properties and porosity levels, especially within the top and bottom layers. Therefore, samples ground to cuboid shape, shown in Figure 8C, were additionally analyzed through density measurements.

Therefore, samples ground to cuboid shape, shown in Figure 8C, were additionally analyzed through density measurements. To achieve the desired cuboid shape, samples were ground on all sides via a 45 micron silicon carbide grinding wheel. Since a small number of samples were prepared in this manner, samples were hand ground by pressing one side of the sample onto the rotating grinding wheel until a plane surface was achieved. This was repeated for all six sides of each sample. The density results of the post-processed parts were shown in Figure 8D. It is interesting to note that the general trend for density measurements first increased and then decreased with an increasing scan speed, which was opposite of the findings for the density of the entire samples. This indicates that the porosity levels in the unstable regions, primarily within the top most layers, significantly affected the average density for the entire sample.

This indicates that the porosity levels in the unstable regions, primarily within the top most layers, significantly affected the average density for the entire sample. The general trend in the porosity analysis here was that increased scan speed results in increased porosity is similar to the results for laser engineered net shaping of 17-4PH stainless steel [32]. It was assumed that the initial powder porosity played a role in the increase observed. An explanation for this result is that increased scan speed resulted in shallower melt pool depths, and decreased amount of time for gasses contained within powder conglomerates to escape before being covered by the following layer.

At higher scan speeds, decreased laser-deposited layer interaction time also produced lower melt pool temperatures and increased viscosity, thus limiting gas removal mechanisms.3.3.3 Effects of Laser Power on Residual PorosityAt high laser powers, the porosity was shown to decrease dramatically, e.g., with porosity as low as 1.2% at a laser power of 485 W. As shown in Figure 9A, the fabricated samples at 275 W showed a wide range of pore sizes distributed throughout the cross-section. With many small pores, a significant number of medium-to-large pores were observed. On the other hand, thin wall samples produced at 485 W, seen in Figure 9B, showed only a few large pores with most pores being very small. Medium sized pores were primarily absent from the printed at 485 W.918Figure 9: Results from laser power study including A. Micrograph of sample cross-section fabricated at 275 W showing a high frequency of large pores, B.

Micrograph of sample cross-section fabricated at 275 W showing a high frequency of large pores, B. Micrograph of sample cross- section produced at 485 W with greatly reduced pore size and frequency, C. Thin wall structure printed at 275 W showing good printability and uniform layers, and D. Thin wall structure exhibiting irregular shape and poor resolution due to a higher laser intensity at 485 W. Both samples were printed using 1000 mm/min scan speed and 2.1 g/min powder flow rate.A possible explanation for this result is that a high laser power as sufficient to maintain a high enough temperature in deposited layers to allow gasses to combine and escape through the surface of the melt pool. In particular, the few trapped large pores in Figure 9B, larger than those seen in Figure 9A, indicated that the increased duration at a high temperature allowed the gasses to combine into larger pores.

It should be noted that samples produced at 485 W exhibited a decreased stability of the melt pool and thus exhibited a part of an irregular shape. Figure 12C shows a thin wall structure fabricated with the laser power 275 W, which is uniform in shape along the build direction due to a relatively good print resolution. In contrast, Figure 9D demonstrates the difficulty of fabricating thin wall structures at a high laser power. The melt pool spread significantly due to decreased melt viscosity at the higher temperatures, meaning that the thin wall structure was much wider than structures produced at lower laser powers. The width of the sample919printed at 485 W was 6.3 mm, while the sample printed at 275 W was 3.8 mm wide. Hence, print resolution was reduced by nearly a factor of two by increasing laser power to 485 W.

Hence, print resolution was reduced by nearly a factor of two by increasing laser power to 485 W. The decreased resolution seen at high laser powers would pose challenges in producing small structures with high dimensional accuracy without reducing laser spot size.3.3.4 Translucency DemonstrationIn order to examine the feasibility of the proposed deposition process in fabricating MAS samples with transparency, or at least translucency, thin wall samples were first prepared using the optimal processing parameters showing minimal porosity. The parameters used included 275 W, 1000 mm/min, and 0.8 g/min for laser power, scan speed, and powder flow rate, respectively. the sample was polished to a thickness of 0.7 mm. Figure 10 shows that a MAS sample with translucency was successfully made through laser direct deposition. The Missouri S&T logo was visible through the sample. It should be noted that thicker samples appeared to be opaque primarily due to cracking and residual porosity.

The Missouri S&T logo was visible through the sample. It should be noted that thicker samples appeared to be opaque primarily due to cracking and residual porosity. Though not showing a transparent sample, this well demonstrated the viability of the proposed laser direct deposition in directly printing MAS samples. In combination with other benefits of 3D printing samples of complex geometries, this process showed great potentials in the fabrication of high strength transparent ceramics. Further studies will be performed in optimizing the processing parameters and their effects on the transparency of laser direct deposited MAS ceramics.Figure 10: Polished MAS sample demonstrating translucency over a grid of Missouri S&T logos920ConclusionsThis paper demonstrated the feasibility of laser direct deposition in rapidly fabricating MAS parts that would be otherwise difficult and expensive to manufacture using traditional methods.

The effects of processing parameters on the microstructure of MAS ceramics fabricated by laser direct deposition was studied, with cracking and porosity being key issues in affecting the transparency of the fabricated parts. It was showed that translucent MAS ceramic samples could be successfully fabricated. The process described in this paper used precursors MgO and Al2O3 for direct laser synthesis, without the need for prior synthesis into MAS powders, typically a lengthy and expensive process. The typical microstructure of MAS ceramics was found to be dominated by columnar grains, orientated in the build direction but at a tilt angle toward the direction of incoming powder flow due to cooling by the powder conveying gas. Microcracks, as a significant defect of MAS ceramics by direct laser deposition, were primarily orientated in the build direction along grain boundaries.

Microcracks, as a significant defect of MAS ceramics by direct laser deposition, were primarily orientated in the build direction along grain boundaries. Future work will include an in-depth study on how grain size, and microcracks are affected by processing parameters including laser scan speed and powder flow rate. It was also shown that increased scan speed resulted in an increase in total porosity. Increased laser power resulted in samples with the lowest porosity levels through increased localized temperatures and longer high temperature durations. However, this reduction in porosity came at the costs of decreased resolution and dimensional accuracy. Future work includes studying the effects of processing parameters on the transparency of laser direct deposited MAS ceramics.AcknowledgementsThe authors would like to thank Dr. Jeremy Watts for helpful discussions, advice, andassistance with sample preparation for microstructural analysis.

The SLE fabricated MAR- M247 samples are investigated using optical microscopy, scanning electron microscopy, x-ray diffraction, and Vickers microhardness measurements. The results show that the average Vickers microhardness values do not show any significant variation with changes in SLE processing parameters. However, the microhardness values are unevenly distributed and show variations along the build direction and the laser movement direction. Overall the hardness values are within ±2 limits for all the SLE deposited MAR-M247 samples. The effect of heat treatment on the microhardness and the residual stress is also investigated. The results show that the microhardness increases and the residual stress decreases after the heat treatment.

The results show that the microhardness increases and the residual stress decreases after the heat treatment. This work is sponsored by the Office of Naval Research through grant N00014-14-1-0658.Keywords: Nickel-Base Superalloy, MAR-M247, Additive Manufacturing, Scanning Laser Epitaxy (SLE), Microstructure, Vickers Microhardness, Residual StressIntroductionNickel-base superalloys are extensively used in the fabrication of the gas turbine hot-section components as they exhibit exceptional high-temperature strength, creep resistance and corrosion resistance. However, the hot-section components have a limited operating life due to material loss from abrasion, oxidation, and corrosion. Currently, these components are scrapped and replaced when the material loss exceeds certain limits. Hence, there is a great interest in developing an additive manufacturing (AM) technology that can restore the parent metallurgy at the damage locations for these parts.

Hence, there is a great interest in developing an additive manufacturing (AM) technology that can restore the parent metallurgy at the damage locations for these parts. A detailed review on AM of nickel-base superalloys may be found elsewhere and is skipped here for brevity [1].Scanning laser epitaxy (SLE) is a laser powder bed fusion (LPBF)-based AM process that creates polycrystalline or equiaxed (EQ), directionally solidified (DS), and single-crystal (SX) structures through the melting of superalloy powders onto similar chemistry superalloy substrates. The galvanometer-controlled movement of the high-power laser (0.2 to 2 m/s) and high-resolution raster scanning (20 to 200 m line spacing) result in superior thermal control over the solidification process.

SLE has previously been demonstrated to successfully deposit a number of popular hot- section superalloys such as IN100 [2], René 80 [3], MAR-M247 [4], René 142 [5], CMSX-4® [6- 8], and René N5 [9, 10] on both similar and dissimilar chemistry substrates.167However, in SLE as the laser performs raster scans atop the powder bed, the samples are subjected to rapid heating and fast cooling resulting in an accumulation of residual thermal stresses in the samples. Characterization of the residual stress in the SLE fabricated parts is crucial, as the residual stress has critical influence on the mechanical properties [11]. Both destructive and non- destructive characterization techniques are available to measure residual stresses. The most commonly used non-destructive methods are X-ray and neutron diffraction, which are capable of providing near-surface and volumetric residual stress measurements in crystalline materials, respectively.

However, these methods are complicated, expensive or time consuming and has drawbacks with regards to textured samples. Destructive methods include hole drilling, sectioning or contouring, and crack compliance or slitting. While destructive methods are capable of providing location-specific residual stress information, they also require the use of additional analysis to determine residual stresses from deformation behaviour [11].MAR-M247 is a nickel-base superalloy that was developed in the 1970s [4]. The alloy demonstrates high creep strength, good castability, and excellent oxidation and corrosion resistance at elevated temperatures. The alloy contains a high volume fraction (62 vol. %) of secondary phase precipitates [Ni3 (Al, Ti)] and a significant amount (13.7 wt. %) of refractory elements such as tantalum (Ta), tungsten (W), and molybdenum (Mo). MAR-M247 is extremely difficult to weld, and to date, limited literature is available on AM processing of MAR-M247 [9].

MAR-M247 is extremely difficult to weld, and to date, limited literature is available on AM processing of MAR-M247 [9]. In the present work, multiple samples of MAR-M247 were produced by altering the processing parameters such as the laser power and the scan speed. These parameters were combined to define a new measure such as the scan energy density (E). The SLE processed samples were investigated using optical microscopy (OM), scanning electron microscopy (SEM), energy dispersive x-ray spectroscopy (EDS), and Vickers microhardness measurements. A previously developed MATLAB code was applied to extract necessary information from the digital images using quantitative metallography techniques [7]. Multiple Vickers hardness measurements were performed along the length and width of the samples. Residual stress measurements were also performed in the deposit region using x-ray diffraction (XRD).

Multiple Vickers hardness measurements were performed along the length and width of the samples. Residual stress measurements were also performed in the deposit region using x-ray diffraction (XRD). The SLE fabricated MAR-M247 samples were subjected to a commercially available heat treatment process and the residual stresses were evaluated for the heat-treated samples as well.Experimental procedure, materials, and methodsA single-mode 1kW Ytterbium fiber laser (IPG Photonics, model: YLS-1000) was used in the present study. A 3-axis galvanometer scanning system (Cambridge Technology) was used to direct the laser on top of the powder bed and focus the laser beam to a spot size of 20 µm Gaussian beam radius. This ensured tight control of the energy density inside the melt pool. The SLE process was conducted inside a process chamber that was purged to less than 10 parts-per-million (PPM) oxygen at atmospheric pressure using 99.999% high-purity Argon.

The SLE process was conducted inside a process chamber that was purged to less than 10 parts-per-million (PPM) oxygen at atmospheric pressure using 99.999% high-purity Argon. No preheating was done prior to the SLE processing and the substrate coupons as well as the pre-placed powder were initially at ambient temperature. The process was carried out on rectangular EQ cast MAR-M247 coupons having dimensions of 31.50 mm x 5.84 mm x 2.54 mm. Each substrate was placed into a 31.50 mm x 5.84 mm recession cut into an IN625 base-plate. Figs. 1(a) and 1(b) show an overview of equipment associated with the SLE process and the placement of substrate coupons and powder- bed, respectively. A raster scan was employed in the current study. The details about the experimental procedure can be found elsewhere and is skipped here for brevity [2].168(a)(b)Figure 1. (a) The SLE workstation, (b) Placement of the superalloy coupons in the SLE process.

(a) The SLE workstation, (b) Placement of the superalloy coupons in the SLE process. The rightmost position shows the blank base to simulate the hollow passages below the tip in the high-pressure turbine blades. Positions 3 and 4 show the bare substrate coupons and positions 1 and 2 show the substrate coupons with pre-placed powder on top.The MAR-M247 powder used in this study was an argon gas-atomized product made by Praxair Surface Technologies. The composition of the powder is reported in Table 1. The morphology, size distribution, and cross-sections of the powder were analysed using OM and SEM. The MAR-M247 powder particles were mostly spherical as shown in Fig. 2(a). The powder was mounted in Bakelite and polished to a mirror finish. The cross-section was analysed under an optical microscope for the inspection of internal porosity, and limited porosity was detected as shown in Fig. 2(b).Table 1: Composition of the MAR-M247 powder (wt.

2(b).Table 1: Composition of the MAR-M247 powder (wt. %)Zr Ni MAR-M247 8.0 10.0 0.7 10.0 6.0 1.0 3.0 1.0 0.015 0.15 0.05 BalCr Co Mo W Al Ti Ta HfBCA Buehler automated saw was used to section the samples for microstructural investigation. The samples were mounted in Bakelite and polished to a mirror finish; starting with 80 grit paper and progressively increasing the size to 1200 grit. The samples were then polished using 5 µm and 3 µm diamond solutions. Finally, the samples were polished using a 0.5 µm colloidal alumina suspension. The polished samples were then etched using Marble’s reagent (50 ml HCl, 50 ml H2O, and 10.0 gm CuSO4) to eliminate the  phases and reveal the microstructure. Imaging was then completed using a Leica DM6000 optical microscope at 100x, 500x, and 1000x magnifications. SEM investigations were carried out on a Hitachi SU8230 SEM. Microhardness measurements were carried out using a Buehler microhardness indenter with a fixed load at 2000 gf.

SEM investigations were carried out on a Hitachi SU8230 SEM. Microhardness measurements were carried out using a Buehler microhardness indenter with a fixed load at 2000 gf. The hardness values were obtained in Vickers microhardness scale.In order to relieve the residual stresses and enable precipitation of the strengthening phases, the samples were heated and kept at 1080 oC for an hour. Thereafter, the samples were cooled at1695.6 oC /minute or faster to a temperature below 870 oC. Samples were then given a typical commercial low-temperature precipitation heat-treat cycle consisting of 12 hours at 870 oC, followed by air cooling at minimum average cooling rate of 5.6 oC / minute to 650 oC to fully develop the strengthening phases.50 m(b)(a)200 mFigure 2. (a) SEM image of the MAR-M247 powder and (b) OM image of the powder cross- section.Results and discussionDeposit characteristicsFigs.

(a) SEM image of the MAR-M247 powder and (b) OM image of the powder cross- section.Results and discussionDeposit characteristicsFigs. 3(a) and 3(b) show two representative OM images of the SLE deposited MAR-M247 on EQ cast MAR-M247 substrates. These two samples were fabricated using E values of 26.51 J/mm3 and 26.84 J/mm3, respectively. Fig. 3(c) shows the OM image of Sample 1 after heat treatment. The deposits were crack-free and dense. An excellent metallurgical bond formed between the substrate and the deposit region, as shown in Figs. 3(a) and 3(b). The deposit region also shows a finer microstructure compared to the substrate region. The microstructural refinement may be attributed to the localized heating and rapid cooling due to the fast moving laser heat source. For all samples, SLE is demonstrated to deposit more than 1500 m of average deposit height in a single pass.The laser power (P) and the scan speed (VS) were varied in the present study.

For all samples, SLE is demonstrated to deposit more than 1500 m of average deposit height in a single pass.The laser power (P) and the scan speed (VS) were varied in the present study. All samples were run with a scan spacing (SS) of 25.4 µm to account for the smaller beam diameter of the fiber laser. The powder layer thickness (tp) was kept as 1650 mm. A new measure was defined based upon theSLE processing parameters – the scan energy density,.SSPPStVEFig. 4(a) illustrates arepresentative microstructure of the SLE deposited MAR-M247 showing columnar dendrites. The columnar dendrites grew in the [001] direction and changed its structure to polycrystalline near170the top surface. The statistical analysis on the PDAS of the columnar dendrites was investigated for each sample in the longitudinal sections. A representative measurement is shown in Fig. 4(a). For each sample, 20 measurements were performed and the average PDAS values were calculated. Fig.

A representative measurement is shown in Fig. 4(a). For each sample, 20 measurements were performed and the average PDAS values were calculated. Fig. 4(b) shows the effect of E on the PDAS. For the SLE deposited MAR-M247, the PDAS increased with E following a power-law relationship as shown in Fig. 4(b). Fig. 4(c) shows the effect E on the average deposit height. With an increase in E, the melt pool extended laterally and transversely allowing more powder to fall inside the melt pool, and thereby, increasing the deposit height.Leading EdgeYSample 1Sample 2Deposit(a)(b)Sample 1 –Heat-Treated(c)SubstrateIIIIIIIIIFigure 3. (a)-(b) Representative OM images of as-deposited MAR-M247 samples showing crack- free and dense deposit. The red arrow represents the build direction.

The deposit region showed hardness values ranging from 460-470 HV for the MAR-M247 sample. The deposit region showed hardness values in the range of 470-490 HV after the heat treatment. Increase in hardness values after the heat treatment is due to the homogenization of the microstructure [4].171However, variations in hardness value were observed in different regions of the deposit. The measurements were performed at 5 and 10 different locations along the Y direction and the X direction, respectively. Figs. 5(a) and 5(b) show the measurement grid points and two different representative indentation marks in the deposit region, respectively. The measurement pattern was chosen to accurately reflect the variation of hardness along the build direction and the laser movement direction.

The measurement pattern was chosen to accurately reflect the variation of hardness along the build direction and the laser movement direction. The average values of the hardness parameters were calculated as follows:(1) /50VH VH101i51jAverage (2) 5/VH VH....,10 3, 2, 1,i51j10,.....,3,2,1Xi (3) 10/VH VH....,5 3, 2, 1,j101i5 ...., 3, 2, 1,jY100 mjVHY1VHY5iVHX1VHX10Figure 5. (a) Schematic of the measurement locations in the deposit and (b) representative indentation marks in the MAR-M247 deposit region.350400450500550Vx1Vx3Vx5Vx7Vx9Vickers Hardness (HV)Location(a)(b)E = 26.84 J/mm^3E = 23.86 J/mm^3VHAVG(HV)E = 19.52 J/mm^335040045050055012345E(J/mm3)(c)VHY1 VHY2 VHY3 VHY4 VHY5VHX1 VHX3 VHX5 VHX7 VHX9350400450500550Vy1Vy2Vy3Vy4Vy5y = 0.1153x + 461.76R² = 0.0006350400450500550182022242628Figure 6.

Variation of Vickers microhardness along the (a) build direction and (b) the laser movement direction for different values of E for as-deposited samples. (c) Effects of E on the average Vickers hardness in the deposit region for as-deposited samples.172The average value of the Vickers hardness was of the order of 460-490 HV for as-deposited samples and the variation was of order ±2 along the build direction and the laser movement direction as shown in Figs. 6(a) and 6(b), respectively. With change in E, no significant changes were observed in the average hardness values as shown in Fig. 6(c). These results are encouraging as it signifies that changing the SLE processing parameters will have a minimal impact on the hardness distribution in the deposit region with most of the variation originating due to the variations in the local microstructure.

Here θ is the angle between the incident beam and the (hkl) lattice plane, d is the lattice spacing of the (hkl) plane of interest, and λ is the incident x-ray wavelength. It has been shown that there is a clear relationship between the diffraction pattern that is observed when X-rays are diffracted through crystal lattices and the distance between atomic planes (the inter-planar spacing) within the material. By altering the inter-173planar spacing different diffraction patterns will be obtained. The inter-planar spacing of a material that is free from strain will produce a characteristic diffraction pattern for that material. When a material is strained, elongations and contractions are produced within the crystal lattice, which change the inter-planar spacing of the (hkl) lattice planes.This induced change in d will cause a shift in the diffraction pattern.

By precise measurement of this shift, the change in the inter-planar spacing can be evaluated and thus the strain within the material deduced. The most commonly used method for stress determination is the sin2ψ method. A number of XRD measurements are made at different ψ tilts. The inter-planar spacing, or 2q peak position, is measured and plotted. The stress can then be calculated from such a plot by calculating the gradient of the line and with basic knowledge of the elastic properties of the material. This assumes a zero stress at d = dn, where d is the intercept on the y-axis when sin2ψ = 0. More complex solutions exist for non-ideal situations where, for example, ψ splitting occurs (caused by the presence of shear stresses) or there is an inhomogeneous stress state within the material [12].

The strong (200) peak at location I represents a unidirectional growth in the MAR-M247 deposit. Fig. 7(c) illustrates the OM image of location I showing a columnar growth in the deposit region. Fig. 7(b) shows the XRD profile at location II in Fig. 3(a). The strong peak at (200) was not present and the deposit region showed a prominent (111) peak. The OM image of location II is shown in Fig. 7(d). Figs. 7(a) and 7(b) also show a few minor peaks at 34°, 38°, and 58°. These peaks were due to the presence of MC (where M = metal) carbides. MAR- M247 contains about 0.05 wt. % C which resulted in the formation of carbide precipitates in the deposit region. Overall, different regions of the deposit showed different XRD peaks for MAR- M247 samples. Figs. 8(a) through 8(c) show the XRD profiles at locations I, II, and III in Fig.

Overall, different regions of the deposit showed different XRD peaks for MAR- M247 samples. Figs. 8(a) through 8(c) show the XRD profiles at locations I, II, and III in Fig. 3(c) in the deposit region showing the variation in crystal orientation in the deposit region of a representative heat-treated MAR-M247 sample.The average stress was of order 257.5 ± 63.3 MPa for the as-deposited sample while 179.2 ± 33.3 MPa for the heat-treated sample. While heat treatments modify the microstructure and precipitation characteristics [4], such treatments were effective in relieving the residual stresses built within the SLE fabricated MAR-M247 deposit. The residual stresses in the deposit region for the heat-treated sample at locations I and III in Fig. 3(c) were 188 ± 26.2 MPa and 170.3 ± 40.4 MPa, respectively.

The residual stresses in the deposit region for the heat-treated sample at locations I and III in Fig. 3(c) were 188 ± 26.2 MPa and 170.3 ± 40.4 MPa, respectively. The residual stress values were of similar order indicating that the heat- treatment was effective in homogenizing the stresses throughout the deposit region.ConclusionsThe difficulties associated with welding of high  nickel-base superalloys are well established. However, the current study showed excellent results for additive deposition of MAR-M247 using an LPBF-based AM approach. The processing conditions were varied and successful deposits were obtained across a broad range of operation parameters. The processing parameters were conveniently combined to define a new measure such as the scan energy density (E). The average deposit height was found to increase with E.

The processing parameters were conveniently combined to define a new measure such as the scan energy density (E). The average deposit height was found to increase with E. However, no such relationship was observed for the average hardness values.The hardness values showed variations of order ±2 along the build and the laser movement directions. The microstructures of MAR-M247 exhibited recrystallization after the heat-treatment. However, the microstructures became uniform after the heat treatment with approximately 31 % reduction in the residual stress. In the future, detailed investigations will be carried out to compare the experimentally measured residual stresses with finite element based simulations. The effect of heat treatment on the grain texture and distribution will also be analyzed using XRD.AcknowledgementsThis work was sponsored by the Office of Naval Research Cyber-Enabled Manufacturing Systems program through grant N00014-11-1-0670.

The authors would like to thank the School of Materials Science and Engineering, Georgia Institute of Technology for providing machine time to perform the XRD experiments, Prof. Hamid Garmestani, School of Materials Science and Engineering, Georgia Institute of Technology for his valuable insight on analyzing the XRD data,175and Mr. David Tavakoli, PANalytical XRD facilities manager, Materials Characterization Facility, Georgia Institute of Technology for his assistance with performing the XRD experiments.DisclosuresDr. Suman Das is a co-founder of DDM Systems, a start-up company commercializing the SLE technology. Dr. Das and Georgia Tech are entitled to royalties derived from DDM Systems’ sale of products related to the SLE research described in this paper. This study could affect their personal financial status. The terms of this arrangement have been reviewed and approved by Georgia Tech in accordance with its conflict of interest policies.References[1] Basak A, Das S.

The terms of this arrangement have been reviewed and approved by Georgia Tech in accordance with its conflict of interest policies.References[1] Basak A, Das S. Epitaxy and Microstructure Evolution in Metal Additive Manufacturing. Annual Review of Materials Research 2016;46:125-49. [2] Acharya R, Das S. Additive Manufacturing of IN100 Superalloy Through Scanning Laser Epitaxy for Turbine Engine Hot-Section Component Repair: Process Development, Modeling, Microstructural Characterization, and Process Control. Metallurgical and Materials Transactions A 2015:1-12. [3] Acharya R, Bansal R, Gambone JJ, Kaplan MA, Fuchs GE, Rudawski N, et al. Additive Manufacturing and Characterization of René 80 Superalloy Processed Through Scanning Laser Epitaxy for Turbine Engine Hot‐Section Component Repair. Advanced Engineering Materials 2015;17:942-50. [4] Basak A, Das S. Microstructure of nickel-base superalloy MAR-M247 additively manufactured through scanning laser epitaxy (SLE).

Advanced Engineering Materials 2017;19:1600690.176[11] Wang Z, Denlinger E, Michaleris P, Stoica AD, Ma D, Beese AM. Residual stress mapping in Inconel 625 fabricated through additive manufacturing: Method for neutron diffraction thermomechanical model predictions. Materials & Design measurements 2017;113:169-77. [12] Fitzpatrick M, Fry A, Holdway P, Kandil F, Shackleton J, Suominen L. Determination of residual stresses by X-ray diffraction. 2005.to validate177.

aracterization of Mechanically Alloyed Polymer Blends for Selective Laser SinteringJulie P. Martin and Dr. Ronald G. Kander Department of Materials Science and Engineering, Virginia Polytechnic Institute and State University, Blacksburg, VA 24061-0237, USACryogenic mechanical alloying (CMA) is presented in this work as an effective technique for creating materials for selective laser sintering (SLS) applications. CMA offers a solid state method for creating micro-composites consisting of finely dispersed phases, which can then be selectively laser sintered into parts containing co-continuous phases. Particle size and shape, microstructure, and melting characteristics of mechanically alloyed particles are discussed in terms of applications to the SLS process. The investigated using scanning electron characteristics of several model polymer blend systems are microscopy, light scattering particle size analysis, and differential scanning calorimetry.

The investigated using scanning electron characteristics of several model polymer blend systems are microscopy, light scattering particle size analysis, and differential scanning calorimetry. Although only polymer/polymer blend particles are studied here, the CMA process is also a viable technique for creating SLS powders using ceramics or metals.INTRODUCTION AND BACKGROUNDrequire high temperatures, solutions, or compatibilizers. However, a solid state processing technique called cryogenic mechanical alloying (CMA) can be used to blend polymers in powder form without the need for solvents, compatibilizers, or a common processing temperature window. Mechanical alloying (MA) is beneficial for blending immiscible polymers and creating novel morphologies with sub-micron phase domains. Traditional thermal blending techniques such as extrusion and injection molding require the polymers to flow on a macroscopic level.

Traditional thermal blending techniques such as extrusion and injection molding require the polymers to flow on a macroscopic level. However, polymer powders blended via MA can be selectively laser sintered into useful part geometries. This allows much of the refined microstructure created during mechanical alloying to be retained in the final part since no macroscopic flow is required during the laser sintering step1,2.Conventional polymer blendingtechniques such as solution and melt blendingMechanical milling (one component) and mechanical alloying (two or more components) are techniques which were originally developed for solid state processing of metals in the late 1960s. Mechanical alloying is currently widely used in the metals industry for producing composite metal powders with fine microstructures. During mechanical alloying, a ball mill is used to produce alloyed powders.

During mechanical alloying, a ball mill is used to produce alloyed powders. The starting materials (in powder or pellet form) are placed in the ball mill vial with two or more metallic or ceramic balls (Figure 1).Figure 1. Schematic of vibratory ball mill vial and balls3.92In a vibratory ball mill, high energy impacts between the balls and the material occur when the mill’s motor vigorously shakes the vial, trapping material between the balls (and between the balls and the vial walls) with each agitation (Figure 2).As CMA occurs, the particles are repeatedly fractured, deformed, and fused together. This process of repeated fracturing and cold-welding causes a refinement in microstructure with milling time. The result is a two-phase lamellar or plate-like microstructure with an interlamellar distance dependent on processing time4 (Figure 3).

The result is a two-phase lamellar or plate-like microstructure with an interlamellar distance dependent on processing time4 (Figure 3). Other processing parameters which affect the composite microstructure include the energy input, which can be controlled by manipulating the ratio of the total ball mass to the powder mass (charge ratio), milling temperature, ball mill design, and number and size of balls used. The milling temperature can be critical because of its effect on both the material ductility and thermally-aided diffusion across interfaces.Figure collision, extensional flow, and fracture (Region A)3.2. High resultingenergy in weldingball-powder-ball (Region B),Figure 3. Two-phase lamellar microstructure of powder particles produced by mechanical alloying5.Because CMA can combine immiscible materials, there are virtually unlimited numbers of polymeric systems that can be blended via mechanical alloying.

Despite this fact, only a few researchers have applied this technique to blending polymers. To date, the field of mechanically alloyed polymers has been explored to a moderate degree at best. This work aims to investigate CMA as a process for producing desirable materials for the SLS process. Although polymer-polymer systems are studied here, this process can be applied to ceramics and metals as well.The field of CMA polymers is currently lacking a systematic understanding of how variations in processing parameters affect the resultant microstructure of the alloyed powders and post-processed parts. Understanding these relationships are crucial because the microstructure dictates the mechanical properties of the material. Therefore, it is important to understand how mechanical milling and alloying change a system’s microstructure in order to engineer a microstructure based on desired macroscopic properties. This work relationships of mechanically alloyed polymeric micro-composites.

This work relationships of mechanically alloyed polymeric micro-composites. While demonstration systems are studied here, a better understanding of how CMA affects these systems will lead to development of SLS materials. A vibratory ball mill wasinvestigates processing-microstructure93used to produce polymeric blends, and several techniques were used to investigate the effects of the mechanical alloying process on these materials.EXPERIMENTAL PROCEDURECryogenic Mechanical Alloying ProcessA novel laboratory-scale vibratory ball mill was designed and built by a colleague at Virginia Tech2. When operated at cryogenic temperatures, the vial is continuously exposed to a liquid nitrogen bath throughout the milling process. The ball mill can also be operated at ambient temperature by omitting the liquid nitrogen.

The ball mill can also be operated at ambient temperature by omitting the liquid nitrogen. The milling vial and balls (shown schematically in Figure 1) are stainless steel; the vial has an inside diameter of 3 inches, and the diameter of each ball is 0.8125 inches. A schematic of the ball mill is shown in Figure 4.Figure 4. Vibratory ball mill with both ambient and cryogenic capabilities designed at Virginia Tech2.Demonstration SystemsMechanically alloyed micro-composites consisting of DTM Duraform™ Polyamide (nylon 12) were blended with polypropylene (PP), Victrex™ poly(ether ether ketone) (PEEK) and DTM polycarbonate (PC). Blends were produced from resins in powder form in a 50/50 volume percent ratio. The powders were cryogenically mechanically alloyed for 1 hour.

Blends were produced from resins in powder form in a 50/50 volume percent ratio. The powders were cryogenically mechanically alloyed for 1 hour. In addition individual components were cryogenically mechanically milled (CMM) at the same conditions.Differential Scanning Calorimetry (DSC)Thermal analysis was performed on as-received nylon 12 powders as well as the CMA blends. A Perkin Elmer DSC7 was used to investigate crystallization and melt properties of the powders. Powders were heated from room temperature to 225oC at a rate of 10o/minute, held in the melt for five minutes, transition, melting, and recrystallization temperatures were determined.then cooled at a rate of 10o/minute.GlassScanning Electron Microscopy (SEM)An International Scientific Instruments SX-40 scanning electron microscope with an accelerating voltage of 20kV was used to image the topology of CMM powders as well as unprocessed powders.

The powders were sputtered with gold prior to imaging.94Particle Size AnalysisA Horiba LA-700 Laser Scattering Particle Size Distribution Analyzer was used to obtain particle size distributions and average particle sizes for CMA blends, CMM individual components and as- received materials. Powders were disbursed in ethanol for analysis.RESULTS AND DISCUSSIONThere are several important material properties to consider when developing materials for the SLS process. These properties include particle size and shape, particle melting characteristics, and particle microstructure. The techniques used in this ongoing study aim to address each of these issues.Particle Size CharacterizationMedian and mean particle sizes averaged from in Table 1 below. Measurements of distribution breadth such as population standard deviation and dispersity index (second moment/first moment) are also shown along with a 95% range, which indicates the size range within which 95% of the particles fall.

All CMM samples were cryogenically mechanically milled for 1 hour, all CMA samples were cryogenically mechanically alloyed for 1 hour.three runs are shownPopulation Std. Deviation (µm) 40.7 39.98 - 0.76 46.50 - 1.18 41.6 46.58 - 0.57 42.6 - 0.26 39.2 79.40 - 4.44 85.3 - 4.37 40.93 - 0.71 41.7 34.34 - 0.43 32.67 - 0.98 26.97 - 0.58 44.4 83.03 - 1.41 47.7 91.07 - 1.46 27.09 - 0.52 43.6 30.87 - 5.32 44.7 28.26 - 1.16 30.53 - 0.72 44.8 37.60 - 1.17 30.53 - 1.15 43.3 33.48 - 2.86 39.47 - 2.0Median Particle Size (µm)Material Mean (µm)As-Received Nylon 12 As-Received PEEK As-Received PC As-Received PP CMM Nylon 12 CMM PC CMM PP CMA PP/Nylon12 CMA PC/Nylon 12 CMA PEEK/Nylon 12Dispersity Index1.3 1.5 1.2 1.3 1.6 1.2 1.4 1.3 1.6 1.595% of Particles fall within range: (µm)9.87-95.50 0.30-144.0 35.63-150.00 9.49-88.88 2.34-83.27 13.56-172.7 3.51-84.70 3.34-82.36 3.51-138.13 2.18-114.00Table 2.

Particle size distribution results.CMM and CMA generally decreases the overall particle size and increases the dispersity index, although the mean and median values for CMM PC were actually higher than the as-received PC. The lower limits of the 95% range may pose problems for SLS, as extremely small particles may not sinter well due to inefficient packing and relatively large surface area per unit volume. Further analysis of particle distributions is necessary for a better understanding of how CMA affects particle size in relationship to the SLS process.95Particle Shape CharacterizationThe effects of the CMM process can be seen in both the shape and topology of the powder particles (Figures 5-9). When nylon 12 is CMA with another polymer such as PP (Figure 9), the effects of the cold welding and fracturing mechanisms of the CMA process are evident.

When nylon 12 is CMA with another polymer such as PP (Figure 9), the effects of the cold welding and fracturing mechanisms of the CMA process are evident. The layered topology of the CMA blend means that each particle now contains both nylon 12 and PP, but the microstructure of these particles cannot be characterized using SEM. The change in both size distributions and shape have implications for powder handling and packing in the SLS powder bed, as well as in the sintering step itself.Figure 5. As-received nylon 12.Figure 6. Nylon 12 CMM 1 hour.Figure 7. As-received PP.Figure 8. PP CMM 1 hour.Figure 9. Nylon 12/PP CMA 1 hour.96Particle Thermal CharacteristicsThe melting and recrystallization characteristics of SLS materials have implications for laser power and speed necessary to sinter the particles.

In addition, once the particles have been melted by the laser, the temperature/speed at which they recrystallize can affect the thermal stresses present in the finished part. It is desirable for one layer to stay soft during a given laser scan so that when the next layer is added, the powder “sticks” to the previous layer and residual stresses are minimized. Therefore, it is desirable to use a system that has a large difference in melting and recrystallization temperatures to maximize the time between the laser contact and recrystallization. The absolute melting (Tm) and/or softening temperature (Tg) is also important, as it contributes to the laser power necessary for SLS processing. Here, the thermal behavior of the pure nylon 12 component is compared to the thermal behavior of nylon 12 present in the CMA blends. The DSC thermograms presented in Figures 10-13 show the thermal behavior of nylon 12 as well as blends of nylon 12/PP, nylon 12/PC, and nylon 12/PEEK.

The DSC thermograms presented in Figures 10-13 show the thermal behavior of nylon 12 as well as blends of nylon 12/PP, nylon 12/PC, and nylon 12/PEEK. The red line is data collected as the as-received or CMA powder sample is heated; the blue line is the data collected as the sample is cooled from the melt. The endothermic melting transition is present in the heat scans, while the exothermic recrystallization transitions are visible in the cooling scans.Figure 10. As-received nylon 12. Figure 11. Nylon 12/PP CMA 1 hour.Figure 12. Nylon 12/PC CMA 1 hour. Figure 13. Nylon 12/PEEK CMA 1 hour.The CMA blends used in this work demonstrate the thermal behavior of several categories of systems. The nylon 12/PP system is one where both phases co-melt during SLS. In the nylon 12/PC system, the PC=145oC).

The nylon 12/PP system is one where both phases co-melt during SLS. In the nylon 12/PC system, the PC=145oC). In the nylon PC phase is amorphous, but is above its Tg in the melting range of nylon 12 (Tg 12/PEEK system, the PEEK phase melts at a much higher temperature than the nylon 12 phase PEEK=345 oC). During selective laser sintering of the nylon 12/PEEK blend near the Tm of the nylon (Tm9712 phase, the PEEK serves as a reinforcement.

During selective laser sintering of the nylon 12/PEEK blend near the Tm of the nylon (Tm9712 phase, the PEEK serves as a reinforcement. Thermal properties of the nylon as well as the other blend component are summarized in Table 2 below.Nylon 12/PP Blend nylon 185 oC Tm 176 oC Melt onset 187 oC Melt end nylon D Hm 44 J/g Tm component - 161 oC 150 oC Melt onset 166 oC Melt end nylon 147 oC Tc 151 oC Tc onset 147 oC Tc end Tc component 114 oC Tc onset - 118 oC 112 oC Tc end D Hc 28 J/gNylon 12186 oC 182 oC 188 oC 119 J/g- 146 oC 150 oC 143 oC -54 J/gnylonNylon 12/PC Blend 185 oC 174 oC 187 oC 48 J/g - - - 150 oC 153 oC 147 oC - - - 24 J/gNylon 12/PEEK Blend 185 oC 178 oC 187 oC 51 J/g N/A N/A N/A 148 oC 151 oC 144 oC N/A N/A N/A 24 J/gTable 3. Thermal properties of nylon 12 and its CMA blends.These DSC samples were heated to a temperature of 225oC, so the PEEK melting and recrystallization temperatures are not shown.

Thermal properties of nylon 12 and its CMA blends.These DSC samples were heated to a temperature of 225oC, so the PEEK melting and recrystallization temperatures are not shown. The heat of melting (D Hm nylon ) for nylon 12 is approximately half that of the pure component. This is because the blends consist of 50/50 by volume of each component. The onset and end of melting occurs at approximately the same temperature in all the blends. The presence of another phase slightly lowers the onset of melting compared with that of pure nylon 12. The onset and end of crystallization occurs at approximately the same temperature in the pure component and in the blends.

The onset and end of crystallization occurs at approximately the same temperature in the pure component and in the blends. Thus, the SLS processing temperature window (Tm-Tc) remains approximately the same for the CMA blends and pure nylon 12.nylon ) and heat of recrystallization (D HcParticle MicrostructureThe characterization of CMA particle microstructure presents many challenges, and several techniques have been attempted to date with limited results. Atomic force microscopy and transmission electron microscopy have yielded images, but the structure seen is believed to be an artifact of the sample preparation rather than the CMA process. Current efforts are focused on using chemical differences of the two CMA phases to determine the size of the phase domains present in the CMA blend particles.

Current efforts are focused on using chemical differences of the two CMA phases to determine the size of the phase domains present in the CMA blend particles. The microstructure of the CMA powders will be investigated as well as the microstructure of the selectively laser sintered parts so that processing-structure-properties relationships can be establishedACKNOWLEDGMENTSThe authors gratefully acknowledge the support of DTM Corporation in Austin, TX. The authors alsothank Jeff Schultz at Virginia Tech for assistance with the ball mill.981Pan, W.J.D. and Shaw, J., Journal of Applied Polymer Science, 56, 1995, 557-566. 2Schultz, J. P., Kander, R.G., and Suchicital, C.T.A., Solid Freeform Fabrication Proceedings, 1999, 311-317. 3Gilman, P. S. and Benjamin, J. S., Annual Review in Materials Science, 13,1983, 279-300. 4 Farrell, M.P., Kander, R.G., and Aning, A.O., Journal of Materials Synthesis and Processing, 4,1996,151-161. 5Maurice, D.R.

aracterization of Mutli-Material Interfaces in PolyJet Additive ManufacturingIvan Vu*, Lindsey Bass†, Nicholas Meisel†, Bruce Orler§, Christopher B. Williams†, David A. Dillard*Biomedical Engineering and Mechanics Department † Mechanical Engineering Department § Chemistry Department Macromolecules and Interfaces Institute Virginia Tech, Blacksburg, VA 24061AbstractRelatively few engineering devices and structures are monolithic, as combinations of materials are often needed to obtain the necessary functionality, performance, weight, and cost requirements. Progress in additive manufacturing now allows multiple materials and even blends of materials to be produced in a single manufacturing process, opening new opportunities for expeditiously achieving functional and performance targets. Just as interactions at interfaces have long been of interest in the area of adhesive bonding, similar issues need to be addressed for printed composite materials.

Just as interactions at interfaces have long been of interest in the area of adhesive bonding, similar issues need to be addressed for printed composite materials. In this study, a Stratasys PolyJet system was used to produce configurations consisting of a soft acrylic layers (TangoBlackPlus) sandwiched by two stiffer acrylic strips (VeroWhitePlus). Several test methods based on the double cantilever beam specimen, a common experimental approach to characterize adhesive performance, were evaluated to characterize the fracture resistance of the assembled layers. Failures nominally occurred at the interface between the two types of materials. Further testing is providing insights into the effects of print direction, postcuring, and interface architecture on the resulting fracture energies. These studies suggest the opportunities for designing printed interfaces with improved performance and durability for multi-material additive manufacturing products.1.

These studies suggest the opportunities for designing printed interfaces with improved performance and durability for multi-material additive manufacturing products.1. INTRODUCTIONAdditive manufacturing (AM) is a unique alternative to traditional manufacturing techniques that embraces the concept of reducing product development cycles and shortening the time-to-market of new products [1]. AM has become a popular method to produce not only prototypes and molds, but increasingly AM is being used to fabricate components for engineered products for a range of industrial applications. AM production of multi-material components and devices is of considerable interest, as it allows the fabrication of elastomeric hinges, integrated grips, and other products where material changes can achieve enhanced functionality [2], reduced part counts, and streamlined assembly.

The ability to selectively place multiple materials with different mechanical properties offers design freedom for producing final parts with tunable failure locus and resistance to fracture.Even with rapid advances of AM technologies, there are still a number of concerns that need to be addressed. For fabrication with a single material, porosity and anisotropy can be issues due to associated with the layer-by-layer printing process, the curing or solidification process, and other factors. Imperfect or weak bonding between layers can lead to delamination and breakage under stress [3, 4]. Surface finishes are often poor and can introduce porosity, making prototypes susceptible to moisture infiltration [4]. Previous research has confirmed print build orientation has a significant effect on the mechanical properties in uniaxial tension of a number of AM systems [5-8], including those printed with multiple polymers.

Researchers have primarily employed the common dog-bone tensile test specimen in their studies of modulus, strength, and fatigue resistance [5, 7, 9-12]. However, to our knowledge, no literature has addressed the characterization of fracture resistance of these materials for mode I fracture (opening mode; wherein tensile stress occurs normal to the plane of the crack).Recognizing prior studies of strength of monolithic and multi-material systems, this study extends characterization by assessing the fracture behavior and resistance of additively manufactured materials tested in laminated beam specimens. Specifically, trilayer constructions involving two printed polymers are used, which results in specimens resembling adhesively bonded joints.

Specifically, trilayer constructions involving two printed polymers are used, which results in specimens resembling adhesively bonded joints. Several specimen geometries are developed and evaluated, including double cantilever beam (DCB), tapered double cantilever beam (TDCB), and peel configurations, all of which have origins in commonly used fracture tests for adhesive bonded joints and laminated composites. The fracture energy and locus of failure are important characteristics used to evaluate the performance of adhesive bonds, and we believe these properties are also important for AM materials, including those containing interfaces between dissimilar components. Of specific interest is whether these interfaces are inherently weaker than the surrounding bulk materials, resulting in a weakened region that could lead to product failure under quasi-static, sustained, or cycling loading.

This study focuses on measuring the quasi-static behavior or such interfaces, proposing specimens that may be appropriate for characterization of a range of AM systems. Properties obtained could be used to optimize multi-material compatibility and improve printing algorithms, perhaps through alternate patterning or architectures of interface, to result in stronger, tougher, and more durable products. The experimental methods employed in this study are presented in Section 2. Results are presented along with a discussion in Section 3. Closure is offered in Section 4.2. EXPERIMENTAL METHODS2.1. MaterialsThe materials involved in this study are two photopolymers used in the Stratasys PolyJet material jetting AM process (Stratasys, Eden Prairie, MN). In this process, layers of multiple acrylic-based photopolymers are selectively deposited onto a build tray via inkjet printing. The jetted photopolymers are quickly cured by UV lamps that are mounted on the printing block.

The jetted photopolymers are quickly cured by UV lamps that are mounted on the printing block. Models printed by this process are incased in a sacrificial support material to ensure dimensional stability, though it can be easily removed after printing via water jet cleaning.two photopolymers: TangoBlackPlus (TB) and VeroWhitePlus (VW). TB is a black photopolymer that exhibits some rubber-like or elastomeric characteristics, while VW is a white, opaque photopolymer that is significantly stiffer in comparison to the TB. Their respective moduli, tensile strengths, andThe material system of interestistheinterface betweenglass transition temperatures (Tg) are conveyed in Table 1. Several researchers have studied this system in the past. Barclift et al. demonstrated when multiple VW parts are printed on the same tray, the part spacing has a significant effect on the mechanical properties [8]. Moore et al.

Barclift et al. demonstrated when multiple VW parts are printed on the same tray, the part spacing has a significant effect on the mechanical properties [8]. Moore et al. analyzed fatigue strength of the TB/VW interface, indicating that the TB/VW interface failed prematurely in 25% of the dogbone specimens tested [2]. Sugavaneswaran et al. observed improved strength in a pure TB tensile specimen by implementing cylindrical reinforcements of VW throughout its structure [13].TangoBlackPlus (TB) 0.1-0.3 1.8-2.4 -10 Table 1. Table of material properties for TangoBlackPlus and VeroWhitePlus. Young's Modulus and Tg for TangoBlackPlus were not provided by manufacturer [14, 15].Property Young's Modulus (MPa) Tensile Strength (MPa) Tg (°C)VeroWhitePlus (VW) 2000-3000 55-65 52-542.2.

Young's Modulus and Tg for TangoBlackPlus were not provided by manufacturer [14, 15].Property Young's Modulus (MPa) Tensile Strength (MPa) Tg (°C)VeroWhitePlus (VW) 2000-3000 55-65 52-542.2. Test Specimen Configurations and FabricationTo evaluate the fracture resistance of trilayer specimens produced with this material system, several double cantilever beam [16, 17] and T-peel test [18] configurations, consisting of a TB layer sandwiched between VW layers, were designed, printed, and tested, In keeping with relevant literature and specimens for testing adhesive bonds, the VW layers will be referred to as adherends and the TB interlayer will be referred to as the adhesive, though these were in fact all printed in a single AM process via a Stratasys Connex 350 system. Each layer deposited is approximately 0.032 mm thick and for all specimens considered, the total thickness of the TB central layers is kept constant at 0.32 mm.

Each layer deposited is approximately 0.032 mm thick and for all specimens considered, the total thickness of the TB central layers is kept constant at 0.32 mm. The sandwich configuration and TB interlayer thickness were common to all specimens but were distinguished by the VW adherend thicknesses or method used to reinforce the adherend through bonded auxiliary adherends.

These configurations, all of 9.9 mm width, included the following, each of which is illustrated in Figure 1:Aluminum-supported DCB (Al-DCB): 1 mm VW – 0.32 mm TB – 1 mm VW sandwiches bonded to 3.8 mm thick auxiliary aluminum adherends machined with endblocks.Fully-printed DCB (FP-DCB): 4.95 mm VW – 0.32 mm TB – 4.95 mm VW sandwiches with integral endblocks printed during fabrication.Steel-backed DCB (St-DCB): same as the unsupported DCB configuration but VW adherends are backed with 0.759 mm thick steel strips to reduce adherend deformation.TDCB: entire specimen, with 0.32 mm TB interlayer and varying adherend thickness printed in their entirety.T-peel: 0.32 mm VW – 0.32 mm TB – 0.32 mm VW sandwiches tested directly in T-peel configuration.a)b)c)d)e)Figure 1. Illustrations of test configurations and photos of representative specimens during testing.

Illustrations of test configurations and photos of representative specimens during testing. Top to bottom are a) the aluminum-supported (AL-DCB), b) fully-printed (FP-DCB), and c) steel-backed DCB (St- DCB), d) TDCB, and e) T-peel geometries.Expected to minimize printing costs and adherend dissipation, the aluminum-supported specimens were made by bonding printed sandwich strips between machined aluminum adherends. The sandwich strips contained a printed chevron notch in the TB interlayer to facilitate failure initiation within the interlayer. Prior to bonding, the 6061-T6 aluminum adherends were abraded with 120 grit sandpaper using a random orbital sander and cleaned in tap water using an ultrasonic cleaner (Fisher Scientific, Waltham, MA) for 3 minutes. The auxiliary adherends were dried in an oven at 110°C for 1 hour. The printed strips were bonded onto auxiliary aluminum adherends using commercially available LORD 320/322 epoxy adhesive (LORD Corporation. Cary, NC).

The printed strips were bonded onto auxiliary aluminum adherends using commercially available LORD 320/322 epoxy adhesive (LORD Corporation. Cary, NC). A paper spacer was placed in the notch to prevent adhesive from bonding the initial crack. The specimens were clamped and left to cure at room temperature for 24 hours. After 4 hours of curing, excess adhesive was removed with a putty knife to allow visibility of the central layers for crack length measurements.Though using more polymer, the fully-printed DCB specimens minimized ancillary specimen preparation, as they were ready to test without additional bonding procedures following printing. As shown in Figure 1, these specimens exhibited the most deformation during testing. Concerns with viscoelastic and plastic dissipation within the polymeric adherends led to the addition of 22 gauge (0.759 mm) steel strips, bonded to the back of the specimen arms using the bonding method described above.

The steel-backed DCB resulted in somewhat smaller adherend deformations. Fully-printed TDCB specimens were also fabricated using the tapered profile generated according to the geometry factor m given by [16, 19]hham1332While typical adhesive have used a value of 2 for m when designing the TDCB [20, 21], a value of 3.5 is chosen for m in an effort to conserve material, while reducing the effects of dissipation and yielding of the adherends during testing. Though using the most printed polymer, these TDCB specimens resulted in reduced adherend deformation and are often advocated for adhesive testing because of the constant strain energy release rate nature of the specimen.In addition to the four DCB configurations, T-peel tests were conducted to evaluate the peel resistance of the TB/VW interface. The same type of sandwich strips used for the aluminum supported DCB specimens were used in unsupported fashion as T-peel specimens.

The same type of sandwich strips used for the aluminum supported DCB specimens were used in unsupported fashion as T-peel specimens. Variations of this configuration involving checkered patterns and interlayer gradient were also tested. The gradient version is illustrated in Figure 1; in addition to VB adherends and a TB central layer, blended layers of TangoBlackShore60 (TB60) and VeroGrey60 (VG60) were also incorporated.2.3. Build OrientationsAM processes allow for components to be printed in different orientations, offering enhanced design and construction flexibility, but also resulting in nominally identical products with very different properties. In PolyJet, printed primitives are jetted in the direction of the travel of the printing block (X-axis); jetted ‘lines’ (composed of several individual drops) are rastered along the Y-axis to complete a layer. Subsequent layers are then built up in the Z- direction.

Subsequent layers are then built up in the Z- direction. The print resolution varies, with respective resolutions in the x-, y-, and z-directions of 0.043 mm, 0.043 mm (as per manufacturer-stated 600 dpi resolution), and 0.032 mm (when1printing multiple materials). Depending on the direction on any given fabricated item, adjacent material may effectively be co-cured, because of near simultaneous deposition, or can result when liquid monomer is deposited later on a cured layer. Of particular interest in this work is the effect orientation of the interfaces and interlayer resulting from the build direction has on the structural integrity of the fabricated product.Figure 2. Illustration of various print orientations used to fabricate specimens. Left: xzy orientation for A) 3D printed (3DP) sandwich strips, B) and C) FP-DCB, and D) TDCB specimens. Middle: fully-printed DCB specimens oriented A) xyz, B) xzy, and C) yxz.

Left: xzy orientation for A) 3D printed (3DP) sandwich strips, B) and C) FP-DCB, and D) TDCB specimens. Middle: fully-printed DCB specimens oriented A) xyz, B) xzy, and C) yxz. Right: Al-DCB specimens oriented A) xyz, B) xzy, C) yxz.For purposes here, a modified form of the orientation nomenclature given by ASTM52921-13 with reference to the TB interlayer [22] is used. The print orientation for each specimen will be designated by the plane of the interlayer and the direction of load application. For example, the DCB and TDCB specimens shown in the left portion of are printed in the ) indicates the direction of crack propagation; the second orientation, where the first letter ( letter ( ) is the ) is the direction of the specimen and interlayer width; and the last letter ( direction of load application.

Both FP-DCB specimens and the 3-D printed (3DP) strips used for Al-DCB specimens were printed in the -orientations to explore the effect of print orientation on fracture resistance, as shown in the middle and right images of Figure 2. AM orientation, as shown in the right portion of strips, used for T-peel tests, were printed in the Figure 2.xzyyzxzyxxzy,, andyzx2.4. Test ProceduresTypewriter correction fluid was applied to the TB bondline as an aid in visualizing the crack tip and paper rulers were affixed onto all DCB specimens in order to facilitate measurements of the crack length. Mode I tests were conducted using a 5800R Instron (Canton, MA) at a crosshead displacement rate of 1 mm/min, unless otherwise noted.

Mode I tests were conducted using a 5800R Instron (Canton, MA) at a crosshead displacement rate of 1 mm/min, unless otherwise noted. Mode I fracture energies are calculated using corrected beam theory (CBT) [16, 19].Ic3ˆ2PFBaaNGwhere GIc is the applied mode I strain energy release rate; a is the crack length; P and  are the applied load and the end displacement respectively at fracture propagation; F is the large is displacement correction; N is the load block correction; B is the width of the specimen; andzyx2aˆthe crack length correction. The large displacement correction F and the load block correction N are calculated by:212231031alaF322122323591891aalalalN4The crack correction 𝑎̂ is used as a means to correct the crack length for a beam that is not perfectly built in. This correction is calculated by plotting the cube root of the compliance (C/N)1/3 vs. the observed crack length.

This correction is calculated by plotting the cube root of the compliance (C/N)1/3 vs. the observed crack length . Representative results are plotted in Figure 3. Extrapolation of the linear fit of this plot will yield as the negative of the abscissa intercept. The R2 value describes the accuracy of the linear fit in correlation to experimental data and given a value of 0.99, this provides further evidence that the CBT is works well for the calculation of fracture energy.aaˆ1.00.8y = 0.0054x + 0.1253 R² = 0.99793 / 1 ) N C/0.60.4(0.20.050050 Crack length (mm)100150Figure 3. (C/N)1/3 vs.

(C/N)1/3 vs. crack length a for aluminum DCB at 1 mm/minAs with the DCB, the mode I fracture energies for the TDCB are also calculated usingCBT [16].313122343.014ammBEPsIcG5sEwhere described in Section 2.1.is Young’s Modulus of VW andmis the specimen geometry factor previouslyThe T-peel tests were performed at a crosshead displacement rate of 254 mm/min, in accordance to ASTM D1876 in the same Instron load frame. All of the above tests were conducted at room temperature.3. RESULTS AND DISCUSSION3.1. Specimen ComparisonTo evaluate the proposed specimen configurations, samples were printed in the same orientation ( -orientation) and interlayer thickness to assess the effectiveness of each configuration (Al-DCB, FP-DCB, St-DCB, TDCB; Section 2.2 and Figure 2). Tests were conducted within 1-2 days of printing in order to minimize effects of physical aging or other phenomena.

Tests were conducted within 1-2 days of printing in order to minimize effects of physical aging or other phenomena. 1200zyx1000800) 2m / J (600c IG400200020406080100120140160180Crack Length (mm)Al DCBFP DCBSt-DCBTDCBFigure 4 shows the variation of mode I fracture energy for all specimens, where two tests were conducted for each specimen type, with good consistency of the replicates.20012001000800) 2m / J (600c IG400200020406080100120140160180200Crack Length (mm)Al DCBFP DCBSt-DCBTDCBzyxFigure 4. GIc vs. crack length for all DCB configurations in the squares (FP-DCB), green triangles (St-DCB), and purple circles (TDCB).orientation: blue diamonds (AL-DCB), redWhile data from two tests per specimen type are not statically significant, the results indicate that the analysis method demonstrates adequate repeatability.

Optical photographs and SEM images of the fracture surfaces are shown in Figure 5. The Al-DCB configuration results in the lowest GIc of 148 ± 8 J/m2 and the loci of failure is largely interfacial, though the failure meanders back and forth between the interfaces at a spatial scale of several millimeters. The layer boundaries are clearly seen in the Al-DCB configuration in Figure 5a, but overall the fracture surface is slightly smoother in comparison to the fully printed configurations.Al-DCBa)10 mm1 mm400 μmFP-DCBb)10 mm1 mmSt-DCB10 mmc)1 mmTDCB10 mm1 mmd)Figure 5. Failure surface of specimens with corresponding SEM images. Plane of the interlayer is oriented on the xz plane, with the direction of crack propagation aligned along the x axisIn contrast to the rather black and white failure surfaces that will be shown later for specimens prepared with other print orientations, the failure surfaces often appear grey to the naked eye.

Polymer blending of the TB and VW may be occurring, as both VW and TB are deposited on each pass of the printer head, resulting in a potential interphase region that is much tougher than the less blended interface between the TB and VW. In the case of the Al-DCB, it is possible that the benefits of polymer blending, suspected for this print orientation with larger adherend thicknesses, were not achieved given the small dimensional scale of the printed assembly. Although surrounded by a thin layer of support material, it is plausible that as the roller repeatedly runs across the print path to level the deposited layers on such a narrow specimen, deformations, such as vibrations or smearing, compromised the structural integrity of the thin specimen.

Some blending of the VW with the support material may have occurred as the400 μm400 μm400 μmroller flattens the deposited droplets and reduced the properties, though it is not clear how this could have affected what appears to be a near interfacial failure to the interlayer. Increasing the thickness of the surrounding VW material could reduce the likelihood of polymer blending from occurring between the two materials or other anomalous effects that led to the very low fracture resistance of the Al-DCB specimens printed in this orientation. We do feel the very low properties likely arose from printing anomalies rather than the specimen configuration, per se.The fully printed specimens were shown to develop polymer fibrils during testing. Polymer fibrillation could arise from surface imperfections or voids, acting as stress concentrations that subsequently expand under the action of the applied tensile stress to form the fibril structure [23, 24].

It was reported that the TB material has a random degree of porosity [2] which explains the development of polymer fibrils during mode I testing. Such behavior is known to be a major mechanism of increased fracture toughness [23, 24] and this is shown on12001000800) 2m / J (600c IG400200020406080100120140160180Crack Length (mm)Al DCBFP DCBSt-DCBTDCBFigure 4 for all fully-printed specimens. SEM images also show signs of localized plasticity that may have been initiated at the boundaries of defects and voids produced by the printing process. It is believed that polymer blending occurs between the TB and VW, as the failure surface appears grey to the naked eye.

It is believed that polymer blending occurs between the TB and VW, as the failure surface appears grey to the naked eye. This is expected as both VW and TB are deposited on each pass of the printer head, resulting in a potential interphase region that is much tougher than the less blended interface between the TB and VW.The fully printed DCB resulted in higher values of GIc , but it is noted that the adherends experienced significant plastic deformation during testing, thereby dissipating energy and exaggerating the measured GIc values. While bonding steel plates to the adherends was thought to200reduce plastic deformation and dissipation, adding them made little difference. In fact, the steel plate-backed DCB configuration resulted in even larger GIc values compared to the non- reinforced, fully-printed configuration. This may be due to the larger amount of energy being dissipated to deform both the VW and the steel plating during testing.

This may be due to the larger amount of energy being dissipated to deform both the VW and the steel plating during testing. The fully printed TDCB also demonstrated similar trends of exaggerated fracture energies as with the DCBs, though this was not forseen as the tapered height was design to reduce adherend dissipation. It is noted that fabricating the TDCB required the most amount of polymer to print due to the tapered profile of the beam height, leading to much longer print times than the both the 3DP sandwich and DCBs.3.2. Print OrientationTo evaluate the effects of print orientation, FP-DCB specimens were printed in three orientations ( ), as shown in the middle image of Figure 2. The variation of fracture energy against crack length for non-backed, fully-printed DCB specimens from the three print orientations is shown in Figure 6. It is noted that extensive polymer fibrillation occurs for all of orientation, where the applied load during the printed DCBs.

It is noted that extensive polymer fibrillation occurs for all of orientation, where the applied load during the printed DCBs. Specimens printed in the testing is perpendicular to the direction of the print build, are shown to have higher fracture toughness than the other two orientations, where the applied load is in the same direction of the print build. This corresponds with prior research that indicates that the strength of PolyJet materials is poor when the load is applied in the direction of the print build, wherein failure occurs predominantly at the interlayer boundaries [25, 26].yzxzyxxzy,,zyx900xyzxzyyxz800700600) 2m / J (500c I400G300200100020406080 Crack Length (mm)100120140Figure 6. GIc vs.

GIc vs. crack length for the fully-printed DCB configurations based on orientationzyxoriented specimen has a grey color or appearance that is evidence of polymer blending between TB and VW, possibly resulting in an interphase region that is much tougher than the less blended interface. For the specimens built in the orientation, there is a clear distinction between the TB and VW sides of the fracture surface, suggesting a much lower degree of polymer blending as VW and TB were printed on separate passes of the print head. Thus, a TB (VW) layer would have deposited on a cured VW (TB) layer deposited in a previous pass. The SEM images indicated in Figure 7a and Figure 7c demonstrate a discretized failure surface as patches of TB, possibly polymer tufts, are randomly distributed across the failure surface. Such observations of discrete failures are reminiscent of spiral cracking as seen from the work of Dillard et al [27].

Such observations of discrete failures are reminiscent of spiral cracking as seen from the work of Dillard et al [27]. While such discrete failures from their work were seen to initiate from solvent exposure, it is predicted that the voids and defects produced by the printing process may have already weaken the TB enough to experience such failures. The FP DCBs were shown to experience significant plastic deformation during testing, which may have exaggerated the measured fracture energies shown above.Figure 7b indicates that the failure surface of theyzxxzyandyy10 mmx1 mmx400 μma)yzx10 mm1 mm400 μmb)zyxyy10 mmc)xzyx1 mmx400 μmFigure 7. Failure surfaces and SEM images of fully printed DCBs printed in the a) orientation.yzx, b)zyx, and c)xzyThe Al-DCB strip is also printed in the same three orientations, as shown in the right image of Figure 2.

The variation of fracture energy against crack length is shown on Figure 8. It is noted that there are significant differences in the measured fracture energies between thealuminum and fully printed configurations. While the fracture energy of the fully printed DCB is seen to be higher in thezyxorientation, this is not seen for the aluminum configuration.900xyzxzyyxz800700600) 2m / J (500c IG400300200100030507090110 Crack Length (mm)130150170Figure 8. GIc vs. crack length for the Al-DCB configurations based on orientation.zyxorientation appears largely interfacial and lacks the development of fibril structures that occurs for the other two orientations. This may be due to printing errors associated with printing specimens of a smaller length scale in contrast to the fully printed DCBs.

This may be due to printing errors associated with printing specimens of a smaller length scale in contrast to the fully printed DCBs. The failure of Al DCB specimens for the and orientations are shown to alternate from one adherend interface to another, somewhat reminiscent of alternating locus of failures reported for adhesive joints [28-31]. Distinct polymer tufts are left on the failure surface after testing as shown on the top and bottom portions of Figure , and c) 7. Failure surfaces and SEM images of fully printed DCBs printed in the a) directions orientation. and Figure 9.

Failure surfaces and SEM images of fully printed DCBs printed in the a) directions orientation. and Figure 9. While the FP-DCB showed that both theAs mentioned earlier, the failure of the Al-DCB printed in thexzyyzxzyx, b)xzyyzxandxzywere equivalent in fracture toughness, this is not seen in the Al-DCB configuration as the orientation is shown to be the highest in terms of fracture toughness between the other orientations.190yzxyzx10 mma)yzx1 mm400 μm10 mm1 mm400 μmb)zyx10 mm1 mm400 μmc)xzyFigure 9. Failure surfaces and SEM images of Al-DCBs printed in the a)yzx, b)zyx, and c)xzyorientation.3.3. Rate DependencyTo evaluate the rate dependency of the TB/VW interface, additional mode I tests were performed on the Al-DCB configuration at crosshead displacement rates of 0.01, 0.1 and 1 mm/min using the Al-DCB printed in theyzxorientation.

The results are shown in Figure 10.800700600) 2500m / J (400c IG300200100025800700600) 2500m / J (400c IG300200100004510.01 mm/min0.1 mm/min1 mm/min6585 Crack Length (mm)105125a)0.01 mm/min0.1 mm/min2345Nominal Crack Velocity (mm/min)1451651 mm/min67b)yzxFigure 10. Fracture results for 3DP sandwich printed in the orientation and tested at three crosshead displacement rates: a) GIc vs. crack length and b) GIc vs. crack velocity showing clear crack velocity dependence.It is clearly shown that the fracture energy of the TB/VW interface is highly rate dependent, as the fracture energy is seen to increase with increasing displacement rate. The resulting failure surfaces are also dependent on displacement rate and are shown on Figure 11.10 mm1 mm400 μma)10 mm1 mm400 μmb)10 mm1 mm400 μmc)Figure 11.

The resulting failure surfaces are also dependent on displacement rate and are shown on Figure 11.10 mm1 mm400 μma)10 mm1 mm400 μmb)10 mm1 mm400 μmc)Figure 11. Failure surfaces and SEM images of Al DCB printed in the displacement rate of a) 0.01 mm/min, b) 0.1 mm/min, and c) 1 mm/minyzxorientation and tested at theFor the lowest displacement rate of 0.01 mm/min, the failure surface appears largely interfacial to the naked eye. As the displacement rate increases, the failure of the TB/VW interface becomes more cohesive, as the TB is being more evenly distributed across each adherend surface. Such differences to crack propagation and crack path selection was seen in the work of Chen et al., where they observed the effects of T stress for symmetric DCBs tested in both quasi-static and low-speed impact conditions [30].

They demonstrated that DCBs with positive T stress levels had a locus of failure that was dependent on the rate of crack propagation, where interfacial failures were seen under quasi-static rates and cohesive failures were seen for the low impact rates [30]. This was consistent with the earlier findings of Fleck et al., who demonstrated that adhesive layers with a positive T stress displayed a cohesive fracture trajectory [32]. In regards to the current findings above, it is speculated that the Al-DCB specimen may also have a positive level of T stress due to thermal and cure shrinkage of the materials [32]. Hence, similar results to Chen et al. [30] are observed when testing the materials at different displacements rates, though further evidence is needed to substantiate this claim.3.4. Pattern ArchitectureIn addition to the original TB/VW interface, a gradient interface was printed wherein the central layers transition from VW to TB in the center, and back to VW.

Pattern ArchitectureIn addition to the original TB/VW interface, a gradient interface was printed wherein the central layers transition from VW to TB in the center, and back to VW. Note that in this configuration, VeroGrey60 (VG60), a mixture of TB with a majority of VW, is also incorporated in addition to TangoBlack60, which is a mixture of VW with a majority of TB. The test strips were printed in theyzxorientation. The results of the T-peel tests are shown below.2520)15TangoBlackGradientN( d a o L1050050100150 200 Extension (mm)250300Figure 12. Load vs. extension curves from T-peel testsTB/VWGradientAverage Force (N)12.4020.08Energy Release Rate2 (J/m Table 2. T-peel results)25054057It is shown that the pure TB/VW configuration is the weakest in terms of peel. The failure surface shows a pure interfacial failure at the TB/VW interface.

T-peel results)25054057It is shown that the pure TB/VW configuration is the weakest in terms of peel. The failure surface shows a pure interfacial failure at the TB/VW interface. While the exact location of failure is not visually clear, the gradient transition had much better peel performance as it shows a 62% increase in average force and energy release rate when compared to pure TB/VW interface.350Figure 13. Failure surface of T-peel specimens. a) TB/VW and b) gradientThe failure surface for the gradient specimen is largely interfacial, where little material is left on one of the adherends. It is clearly seen that an improvement on peel performance can be accomplished by utilizing different print patterns and geometries at the central layers.4. CONCLUSIONSThis study proposes several test configurations to assess the mode I fracture toughness of polyjet materials.

CONCLUSIONSThis study proposes several test configurations to assess the mode I fracture toughness of polyjet materials. Various DCB and TDCB configurations are constructed and tested to establish which specimen configuration would best represent the fracture toughness. It is demonstrated that all fully printed specimens experience some degree of polymer fibrillation and adherend plasticity during testing, which may have exaggerated the measured fracture energies. The addition of bonded steel plates on the adherend arms were not sufficient to reduce the plasticity effect. The fully printed TDCB demonstrated the same trends of exaggerated fracture energies for unknown reasons, as the tapered profile of the beam height was thought to limit adherend dissipation. It is noted that using the TDCB specimen requires using more VW material to print the tapered profile, leading to longer print times when compared to the other configurations.

It is noted that using the TDCB specimen requires using more VW material to print the tapered profile, leading to longer print times when compared to the other configurations. The 3DP sandwich bonded to Al adherends is thought to be most representative specimen for mode I testing of the polyjet materials as the Al adherends were able to minimize plastic dissipation of the adherends, though it required the longest time to fabricate. It is seen that 3DP sandwich orientation has the lowest fracture toughness, though the reasons specimens built in the behind this are unclear. Given the small dimensional scale of the printed assembly, it is speculated that the roller introduced deformations and smearing of the materials that compromised the structural integrity of the printed assembly. Some blending between the support material and VW may have also occurred and led to the reduced properties and aforementioned failures seen above.

Some blending between the support material and VW may have also occurred and led to the reduced properties and aforementioned failures seen above. It is predicted that increasing the VW boundary would reduce the likelihood of polymer blending from occurring between the two materials or other anomalous effects that would led to the very low fracture resistance of the Al-DCB specimens printed in this orientation. The failure surface of the TB/VW interface demonstrates a random fracture trajectory as TB is randomly dispersed and distributed between the adherend interfaces. SEM images for specimens printed in the orientations indicate an alternating fracture surface that is somewhat reminiscent of alternating failures observe for adhesive jointszyxyzxandxzy[28-31]. After subsequent development of fibrils during testing, polymer tufts of TB are left behind on the failure surface.

After subsequent development of fibrils during testing, polymer tufts of TB are left behind on the failure surface. Signs of localized plasticity are also observed around the polymer tufts, believing to have occurred at the boundary of defects or voids produced by the printing process.The fracture energy is seen to be highly rate dependent, as an increase in displacement rate leads to increased fracture energies. Increasing the displacement rate also leads to failures that transition from purely interfacial to alternating cohesive failures. It is speculated that T stresses may have developed as a result of thermal and cure shrinkage printing. Such rate dependent failures are consistent with previous findings by Chen et al and Fleck et al.

Such rate dependent failures are consistent with previous findings by Chen et al and Fleck et al. [30, 32], though more evidence is needed to substantiate this claim.T-peel tests were performed on the Al-DCB specimens with different pattern architectures implemented at the central layers in order to observe any increased peel performance when compared to the TB/VW interface. It is demonstrated that the gradient pattern had much better peel performance than the TB/VW interface, being 62% higher in terms of both average peel force and energy release rate.While the viscoelastic nature of polyjet materials tend to introduce complexities during deformation, it is shown that these materials can offer adequate fracture performance if the proper print parameters are selected. Other combinations of material and pattern architecture not considered in this study may offer other design options for enhanced fracture toughness.5.

Other combinations of material and pattern architecture not considered in this study may offer other design options for enhanced fracture toughness.5. ACKNOWLEDGMENTSThe authors would like to thank the Biomedical Engineering and Mechanics (BEAM) Department and the Provost’s Office for partial support of IV, the DREAMS Laboratory and the Mechanical Engineering Department the Macromolecules and Interfaces Institute at Virginia Tech for fostering interdisciplinary research in polymer science. We also acknowledge adhesives and bonding advice provided by the LORD Corporation.forfabricatingthe specimenstested, and6. REFERENCES1.2.3.4.5.6.Kęsy, A. and J. Kotliński, Mechanical properties of parts produced by using polymer jetting technology. Archives of civil and mechanical engineering, 2010. 10(3): p. 37-50. Moore, J.P. and C.B. Williams. Fatigue Characterization of 3D Printed Elastomer Material. in International Solid Freeform Fabrication Symposium. 2012.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCharacterization of PLA/Lignin Biocomposites for 3D PrintingJohn Obielodan1, Kevin Vergenz1, and Danyal Aqil1, Joseph Wu2, and Laurel Mc Ellistrem31Mechnical Engineering Department, University of Wisconsin-Platteville, Platteville, WI 538182Chemistry Department, University of Wisconsin-Platteville3Material Science and Engineering Center, University of Wisconsin-Eau ClaireAbstractA greater proportion of polymer-based three-dimensional (3D) printing materials aresynthetic petroleum derivatives that are not biodegradable, contributing to environmental pollution and have potential adverse effects on human health. Polylactic acid (PLA) is currently the most widely used among the bio-based alternatives.

Polylactic acid (PLA) is currently the most widely used among the bio-based alternatives. This work explores alternative environmentally friendly bio-based polymers sourced from sustainable crop and forest biomass derivatives for 3D printing. Various blends of PLA/Organosolv lignin were extruded for fused filament fabrication (FFF) 3D printing process. The processing parameters and results of mechanical and thermal properties of fabricated test specimens of the biocomposite with up to 40wt% lignin concentration are presented. Results indicates that lignin, a low-cost waste product of pulping for the paper industry and bioethanol fuel production could serve as a key component of new biocomposite polymers for 3D printing applications.IntroductionMost utility polymers are synthetic, derived from petroleum resources. Their synthesisinto a great variety of desirable materials compels dependency, absent competing viable alternatives.

Their synthesisinto a great variety of desirable materials compels dependency, absent competing viable alternatives. Their collective drawbacks include processing related carbon footprint addition to the environment, non-biodegradability, and end of life disposal leading to pollution of the ocean waters. Pollution of the ocean waters presently constitute a danger to many marine animals through entanglements and ingestions that results in their deaths and in some cases, hormonal changes that potentially could affect humans that depend on them for foods [1-3].Environmentally friendly and sustainable forest resources generating neutral carbonfootprint have been considered as alternative sources of functional and engineering grade biodegradable polymers. Polylactic acid (PLA), an aliphatic polyester sourced from renewable agricultural products is the most widely used among the bio-based polymers [4].

Polylactic acid (PLA), an aliphatic polyester sourced from renewable agricultural products is the most widely used among the bio-based polymers [4]. It is biodegradable and possesses better room temperature strengths and stiffness compared to acrylonitrile butadiene styrene (ABS), the most widely used material for extrusion-based fused filament fabrication (FFF) 3D printing applications. PLA is easier to print with or without998substrate heating. It is less susceptible to warpages due to non-linear contraction of printed parts in un-enclosed print platforms. It is however brittle, has relatively low glass transition temperature and low ultraviolet light resistance, which are major drawbacks. Its good 3D printability makes it a choice candidate matrix material for natural fibers in the synthesis of biocomposites.Several studies have been conducted on the use of bio-materials as fillers in PLA polymermatrices to achieve differing objectives.

PLA has been demonstrated as a matrix material for natural fibers such as flax, cellulose, kenaf, hemp, bamboo, coir, carbon, ramie, jute, and others [5-11]. Lignin, a product of forest biomass is a subject of active research as a filler material in different synthetic [12-17] and bio-derived thermoplastic [18-21] polymer matrices. Lignin is the glue that binds cellulose fibers together and provide stiffness in plants. It is the second most abundantly available natural polymer after cellulose. As a by-product of pulping for the paper industry, it is yet to attract high value niche applications despite its perceived potentials. Depending on biomass source and processing routes, different strains of lignin are produced, such as alkaline lignin, kraft lignin, organosolv lignin, and others. Each processing route produces lignin macromolecule with peculiar functional groups and properties.

Each processing route produces lignin macromolecule with peculiar functional groups and properties. Most of the earlier work on valorizing lignin through polymer composite synthesis were generic investigations for possible parts manufacturing using injection molding. Recently, different groups have focused on investigating the 3D printability of different lignin-based polymer composites.3D printing, a group of processes in which solid objects are fabricated through layer-wisedeposition of materials has provided a platform for fabricating structures for varieties of applications using different materials. The processes are particularly well-suited for fabricating composite materials, some of which may be difficult to process using traditional methods. In polymer processing, vat photopolymerization 3D printing processes have been investigated for fabricating parts using lignin-based composite materials [22-23].

In polymer processing, vat photopolymerization 3D printing processes have been investigated for fabricating parts using lignin-based composite materials [22-23]. FFF-based 3D printing, being the most prevalent 3D printing process because of affordability is widely used for new materials development. Although, the process is prone to porosities and anisotropic properties of fabricated parts [24]; with continuous machine and process innovations, the qualities of its fabrications would eventually match injection molded structures. FFF 3D printing of lignin-based polymer composites were recently investigated. Eleni Gkartzou et al. [25], investigated the printability and performance evaluation of kraft lignin-filled PLA. The work demonstrated limited success in printing PLA/kraft lignin composites because of poor rheological flow of the melts. Only 5wt% kraft lignin filled PLA was successfully printed. Higher concentrations could not be printed because of rheological challenges.

Only 5wt% kraft lignin filled PLA was successfully printed. Higher concentrations could not be printed because of rheological challenges. Increasing concentrations of kraft lignin in PLA resulted in more brittle composite material. In another work, Ngoc A. Nguyen et al. [26-28] successfully produced high strength blends of up to 40wt% organosolv lignin with acrylonitrile butadiene rubber and reinforcing short carbon fibers in ABS matrix using FFF 3D printing. This synthesis of lignin in petroleum-based materials like ABS would at best be partially biodegradable and would not solve the environmental problems occasioned by process-based carbon foot print and after life polymer material disposals. Although, it provides a route for lignin valorization.999In the present work, the printability and performance of organosolv lignin filled PLAcomposites are explored. The aim is to develop a new utility biodegradable polymer using renewable and biodegradable biomass resources.

The aim is to develop a new utility biodegradable polymer using renewable and biodegradable biomass resources. Different organosolv lignin and PLA blends were synthesized up to 40wt% lignin and printed using FFF-based 3D printing process. Mechanical properties and thermal behaviors of printed samples were evaluated.Experimental ProceduresPulverized PLA made from pellets supplied by NatureWorks LLC was used as matrix forall composite synthesis in this work. The organosolv lignin used were partly sourced from American Science and Technology LLC while some were synthesized in-house. The pulverized PLA had an average diameter of 1.8 mm. The organosolv lignin were received from synthesis as dried solid blocks of materials that were later crushed and sieved to mesh size 12. All PLA and lignin materials were placed in vacuum oven at 45 oC for 24 hours before use. Composite material blends were synthesized by weight percent of constituents.

All PLA and lignin materials were placed in vacuum oven at 45 oC for 24 hours before use. Composite material blends were synthesized by weight percent of constituents. For each composition, material blends were mechanically mixed thoroughly to obtain visually homogenous blends before pouring into a single screw Filabot EX2 extruder. To demonstrate the effects of composite synthesis with the aid of chemical modification of the lignin functional groups for enhanced surface bonding and miscibility with PLA, 20wt% and 30wt% lignin compositions were treated with silane compatibilizer. For these compositions, the lignin contents were pretreated with (3- aminopropyl) triethoxysilane before blending with PLA. The Filabot EX2 extruder used has a (cid:2873) − (cid:1861)(cid:1866)(cid:1855)ℎ outer diameter screw, length to diameter (L/D) ratio of 12, pitch of ½ inches, and(cid:2876) compression ratio of 2.5:1.

The extrusion system setup included the extruder, a Filabot airpath that comprise of a set of six fans that provide forced air-cooling for the single filament being extruded, and a Filabot spooler. A digital measuring device mounted on the Filabot spooler was used for real-time measurement of extruded filament diameter. A 1.75 mm diameter nozzle was used for all filament extrusions. The extrusion temperatures were varied between 160 and 170 oC depending on composition. The extrusion temperature, extrusion speed, airpath cooling fan speed, and spooling speed were adjusted until stable, consistent filament diameter of 1.75 ± 0.05 mm was achieved. Spools obtained from Filabot LLC were used to collect filaments in rolls in sufficient quantities needed for test fabrications. Extruded filaments were stored in vacuumed plastic bags to prevent moisture absorption.Creality CR-10S table top FFF-based 3D printer with Cura software was used for fabricating all tensile specimens.

Tensile tests were conducted according to ASTM D638. Test specimens were fabricated with five replicates for the following weight percent of organosolv lignin in PLA: 1, 5, 10, 15, 20, 25, 30 and 40. The machine was set up for 100% density. All fabrications were made using 0.4 mm 3D printing nozzle size. The nozzle temperature was varied between 200 oC and 205 oC while the bed temperature was maintained at 60 oC. A 50 kN MTS tensile testing machine was used.Fractographic images of representative fractured tensile samples were obtained usingHitachi SU1510 scanning electron microscope (SEM) on the secondary electron mode.1000Thermal behavior was evaluated for composite material samples using differential scanning calorimetry (DSC) Q20 V24 made by TA instruments. DSC samples were subjected to three heating circles. The first was to 110 oC at 10 oC /min and then cooling to 25 oC to eliminate moisture. The second and third were to 200 oC at 10 oC /min.

DSC samples were subjected to three heating circles. The first was to 110 oC at 10 oC /min and then cooling to 25 oC to eliminate moisture. The second and third were to 200 oC at 10 oC /min. The tests were used to collect information on glass transition temperature (Tg), cold crystallization temperature (Tcc), melting temperature (Tm), enthalpy of cold crystallization (∆Hcc), and enthalpy of melting (∆Hm).Results and DiscussionFilaments with compositions up to 40wt% unmodified organosolv lignin in PLA wereextruded successfully. All the melts in the current work demonstrated good rheological properties that enabled material flow and sustained filament formation. This is in contrast to the maximum of 5wt% kraft lignin concentration previously reported [25]. Extruded filaments were used for 3D printing tensile specimens with mechanical properties shown in Fig. 1 to Fig. 3. Similar to earlier reports with any type of lignin, the tensile strengths shown in Fig.

1 to Fig. 3. Similar to earlier reports with any type of lignin, the tensile strengths shown in Fig. 1 decreased with increasing weight fractions of organosolv lignin. This trend is due to limited interfacial bonding between the constituents, especially because of the hydroxyl groups present in unmodified lignin macromolecules. Weight transfer between unmodified lignin and the PLA matrix material is limited. Increase in miscibility is realizable by substitution of the hydroxyl groups with other functional groups that could enhance bonding between the matrix and filler materials. Unlike kraft lignin-filled PLA that became increasingly more brittle than neat PLA at higher concentrations of the filler material, organosolv lignin filled PLA became increasingly more ductile. Strength increases due to lignin pretreatment with silane is demonstrated in Fig.1 for 20wt% and 30wt% compositions. At 20wt% composition, strength increased from 34.4 MPa to 38.2 MPa, an 11% increase.

At 20wt% composition, strength increased from 34.4 MPa to 38.2 MPa, an 11% increase. Composites with up to 25wt% unmodified lignin had a minimum of 32.8 MPa tensile strength, which is significantly higher than the 24 MPa recorded for ABSPlus. ABSPlus is one of the common ABS materials for FFF 3D printing marketed by Stratasys, the industry leader in extrusion-based 3D printing industry. The tensile strength of ABSPlus is slightly higher than those for 30wt% and 40wt% unmodified lignin filled PLA. However, with modification using silane, the 30wt% lignin composite recorded higher strength. 40wt% lignin composition was not modified in the present work.Figure 2 shows stiffnesses of composites with unmodified lignin ranging from 2.45 GPafor 30wt% lignin composition to 3.15 GPa for 10wt% lignin composition. Those with silane treatment were consistently high at 3.18 GPa and 3.13 GPa for 20wt% and 30wt% organosolv lignin.

Those with silane treatment were consistently high at 3.18 GPa and 3.13 GPa for 20wt% and 30wt% organosolv lignin. As can be seen, the stiffnesses of the PLA/lignin composites and neat PLA are not significantly different.One of the limitations of PLA is its brittleness. As shown in Fig. 3, neat PLA at 4.24percent elongation (%EL) is much more brittle compared to composites with higher organosolv lignin composition. Organosolv lignin addition helps to plasticize PLA. However, at higher filler material loading, higher variabilities of the percent elongation results. The higher compositions also demonstrate much higher percent elongation compared to ABSPlus at 6.85. This is an indication of superior toughness over ABS.

The fractographs in Fig. 4a to 4c show well bonded internal structures with few incidences of small air pockets that are typical features of FFF-based 3D printed parts. Figures 4a and 4b for 10wt% and 15wt% lignin composition respectively, show more incidences of air pockets than the 20wt% lignin sample in Figure 4c. This could be because the 20wt% is a little more ductile and could have been less viscous during deposition enabling better material infill. There is no indication of lack of homogeneity in the compositions of the thermoplastic materials.(a)(b)(c)Figure 4: Representative SEM fractographs of tensile specimens for (a) 10wt%, (b) 15wt%, and (c) 20wt% lignin composition1003Thermal AnalysisFigure 5 shows thermograms obtained from DSC tests of the composite samples while Table 1 shows extracted thermal parameter data. The glass transition temperatures are slightly higher at lower lignin contents.

The glass transition temperatures are slightly higher at lower lignin contents. Each sample recorded a single glass transition temperature and melt temperature. Ordinarily, this could be assumed to be indicative of high miscibility of the components, but this assumption is unsupported by the decline in tensile strengths (Fig.

The single melting temperature for each of the samples could be indicative of the formation of a single type of crystal during cold crystallization. For all samples, the differences in melting temperatures are not significant. According to the DSC data, The glass transition, cold crystallization, and melting temperatures are virtually unaffected by lignin content. Although compositions with chemically modified lignin yielded higher strengths and stiffnesses, their thermal properties are similar to those with unmodified lignin.ConclusionThis work has demonstrated the viability of organosolv lignin-filled PLA for meltextrusion at high filler material loading. This is due to favorable rheological flow of the composite melt that support sustained filament formation. The composite filament is printable for all compositions using 0.4 mm diameter 3D printing nozzle. This is possible because the composites can endure high shear rate prevalent in FFF-based 3D printing process.

This is possible because the composites can endure high shear rate prevalent in FFF-based 3D printing process. Although filament extrusion and 3D printing of the materials are not a challenge, the low miscibility of the constituents with unmodified lignin could limit organosolv lignin weight fractions in PLA for load bearing applications. Based on the results of mechanical tests obtained, compositions up to 25wt% organosolv lignin without modification out performs ABS material for room temperature applications. With modifications, higher weight fractions of organosolv lignin would performs better than ABS.Acknowlegment: This research was supported by WiSys and UW System applied research funding programs (ARG, AR-WiTAG, PDF). The views expressed herein are those of the authors and are not necessarily those of WiSys or UW SystemReferences[1]Haward, M., 2018. Plastic pollution of the world’s seas and oceans as a contemporary challenge in ocean governance.

In this contribution, powder requirements for SLS, structure-property relationships and appropriate methods for powder characterization are reviewed. Effects of particle size, particle shape and surface functionalization on flowability, packing density and tribo-charging will be discussed for commercial PA12 laser sintering powders (virgin vs. aged), polyolefin and polyester powders. The possibilities of dry particle coating as an efficient method to tailor powder flowability, bulk density and charging behavior are demonstrated. The capabilities of a Schulze ring shear tester, a powder tensile strength tester, a thermally controllable ring shear apparatus and a model experiment mimicking the powder spreading are discussed to assess SLS processability.IntroductionPowder bed fusion processes (PBF), such as e.g.

selective laser sintering (SLS) / laser sintering (LS), selective laser melting (SLM) or selective electron beam melting (SEBM), allow for manufacture of complex functional parts without the need for any tools or molds. Briefly, in these processes the part is built layer by layer from powders, which are selectively melted. In SLS, a thermoplastic powder –preferentially a semi- crystalline polymer, mostly polyamide 12 (PA12)- is selectively melted by a laser (typically CO2, 10.6 μm wavelength): The manufacture is performed in SLS machines in a building chamber, which is typically heated to a process temperature slightly below the melting temperature of the respective polymer feed powder –in case of PA12 around 167 °C. First, the powder material is applied at a certain powder layer thickness (typically 100 μm to 150 μm) to the building platform by means of a spreading unit, which frequently is either a rake or a roller coater.

Then, the applied powder is selectively melted, where the later (solid) part is desired. Finally the building platform is lowered by around one powder layer thickness and a new building cycle starts with the powder spreading step. From this short sketch of the SLS process it already becomes obvious, that there are many factors that will affect processability and part quality in terms of e.g. dimensional accuracy and mechanical properties. Factors that influence SLS processability have been intensively discussed in literature (c.f. e.g. [1-5]) and are summarized in the scheme depicted in Figure 1. One can subdivide factors that determine SLS processability into intrinsic material properties and (extrinsic) bulk solid properties.

[1-5]) and are summarized in the scheme depicted in Figure 1. One can subdivide factors that determine SLS processability into intrinsic material properties and (extrinsic) bulk solid properties. Intrinsic material properties, like optical, thermal (melting temperature, solidification temperature, degree of crystallinity, heat capacity, thermal expansion coefficient) or rheological properties (melt viscosity) depend on the formulation of the plastic, i.e. type of polymer and its structure, the molar mass distribution, cross-linking, additive enhancement. The given material characteristics influence for example the so-called sintering window, which can be employed during PBF of polymers, i.e. the temperature difference between onset of melting and onset of solidification temperature or the melting and melt coalescence behavior and, thus, are e.g. important to identify proper building chamber temperature or energy input during laser illumination [6].

important to identify proper building chamber temperature or energy input during laser illumination [6]. The extrinsic properties of the polymer powder, i.e. its bulk solid properties like flowability, packing characteristics (c.f. bulk vs. tapped density) depend mainly on inter particulate interactions, which are determined for example by the particle size distribution, the particle shape or the particle surface characteristics (surface roughness, surface chemistry). The mentioned particle characteristics are a function on the processes applied for particle formation and / or functionalization.779Within this account we will review techniques that allow for characterization of extrinsic SLS material properties, i.e. bulk solid and particle characteristics.

• Melt rheologySLS processabilityFigure 1: Overview on key characteristics of thermoplast powders influencing processability in selective laser sintering (SLS).Bulk solid characteristics of commercial SLS powdersMost of the currently employed SLS powder by market share (> 90 %) is made up of polyamides (PA), mainly polyamide 12 (PA12) besides some PA11 and PA6. The PA12 materials are available as a standard grade but also as special grades e.g. enhanced with flame retardants or as glass, carbon or aluminum filled systems. PA particle systems for SLS typically are produced either by direct polymerization (c.f. Orgasol grades by Arkema) or precipitation (c.f. e.g. PA2200 (EOS) or DuraForm PA (3D Systems)). Besides PA powders, there are commercial SLS powders made up of polystyrene (PS), polypropylene (PP), thermoplastic eleastomers (TPE) or polyether ether ketone (PEEK) available [4-9].

Besides PA powders, there are commercial SLS powders made up of polystyrene (PS), polypropylene (PP), thermoplastic eleastomers (TPE) or polyether ether ketone (PEEK) available [4-9]. For further information on commercial SLS particle systems and recent developments, please refer e.g. to [8] or [4]. SLS polymer powders also can be obtained by cryogenic grinding (c.f. PA1101 (EOS)), however, there can be issues with inferior flowability due to edged particle shapes being typical for comminution products [10] and, in consequence, inferior processability [11]. Commercial PA12 SLS powders of good processability typically have particle sizes between 45 μm and 90 μm. Moreover, they are often surface-functionalized with flowing aids (e.g. pyrogenic silica or alumina), which increases surface roughness and allows for minimization of particle interactions and, thus, improvement of flowability and packing density [12-18].

pyrogenic silica or alumina), which increases surface roughness and allows for minimization of particle interactions and, thus, improvement of flowability and packing density [12-18]. Figure 2 exemplarily depicts SEM pictures of PA2200 (EOS), a PA12 SLS powder. The particle shape (Figure 2, left), which is often referred to as ‘potato-shape’, is typical for polymer particles produced by precipitation, more precisely liquid-liquid phase separation and subsequent crystallization (for details on this process, please refer to our recent work [18]).780Figure 2: Scanning electron microscopic (SEM) pictures of PA12 SLS powder PA2200 (EOS). Left: ‘potato- shaped’ particles being typical for precipitated particles. Right: particle surface with flowing aid.The surface functionalization of PA2200 with flowing aid becomes obvious from the material contrast being visible in Figure 2, right.

Right: particle surface with flowing aid.The surface functionalization of PA2200 with flowing aid becomes obvious from the material contrast being visible in Figure 2, right. Furthermore, the PA2200 particles are characterized by a rather narrow particle size distribution with x10,3 = 45.8 μm, x50,3 = 61.4 μm, x90,3 = 81.0 μm and a span ((x10,3 - x90,3)/ x50,3) of 0.57. Narrow particle size distributions and the absence of fines is advantageous with respect to flowability and dense packing.

Narrow particle size distributions and the absence of fines is advantageous with respect to flowability and dense packing. To summarize, the key to optimization of PBF processability concerning bulk solid properties is control of(i) particle shape (rounded particle shapes are advantageous) (ii) particle size distribution (narrow distribution is advantageous) (iii) particle surface roughness (rough particles allow for reduction of inter particulate forces and, thus,better flowability).Strategies to achieve these goals will be discussed in the following section in line with methods for characterization of powder flowability and spreadability.Methods for characterization of powder flowabilityThe determination of bulk solid properties at ambient temperature frequently already allows for a first assessment of the processability of the PBF material under process conditions.

Powders of good flowability are essential, they typically show good spreadability during the powder deposition step and, thus, guarantee a dense and homogeneous powder layer and, consequently, parts of low porosity [9, 19, 20]. Various methods have been proposed so far, although there is no consensus, which method is most appropriate with respect to mimicking the stress conditions during powder spreading and, thus, reliable prediction of PBF processability.Hausner ratio and Carr indexFrequently, as a rough estimate of flowability, the so-called Hausner ratio (HR) or the Carr index (C) are applied, see equations 1 and 2.

Table 1). Moreover, bulk density and tapped density highly depend e.g. on the mode of filling a measurement cylinder for determination of the volume occupied by the bulk solid, or the mode of consolidation during tapping (tapping frequency, stroke, number of taps etc.) [21]. Thus, the obtained results may show a rather large variance, especially if no automated tapper devices are employed, i.e. if the powder consolidation instead is done(cid:1829) = 100 (cid:942)781(1)(2)manually. However, these methods are among the few standardized flowability measurement procedures, which fuels their widespread application, despite their flaws.Shear testersShear testers, like e.g. a ring shear cell, allow for the determination of powder flowability under well- defined conditions [16, 18, 22] and, thus, the comparison of the flowability of different materials via the flow function ffc.

in [16]. An assignment of flowability, flow function ffc (and Hausner ratio, HR) is given in Table 1. Figure 3 exemplifies flow functions determined for PA11 PBF materials at ambient temperature and demonstrates how flowing aids and the removal of fines help to improve powder flowability.(cid:1858)(cid:1858)(cid:1855) =a P/1250ffc = 1ffc = 2ffc = 4cσ h t g n e r t s1000750ld e y d e n i f n o c n ui5002500■PA11 (as precipitated) PA11 (classified & dry coated with flowing aid) PA1101 (EOS)ffc = 104000 consolidation stress σ Figure 3: Flowability of PA11 powders as determined by shear experiments using a Schulze ring shear tester: narrowing of particle size distribution and dry coating with flowing aids allow to tune flow behavior from cohesive (2 < ffc < 4, red open symbols) to easy flowing (ffc > 4, red filled squares).

The flow function ffc of commercial PA1101 (EOS) PBF material is given for reference.0100020003000500060001 / PaThe red open symbols depict the flow function of a PA11 powder obtained by liquid-liquid phase separation and (isothermal) precipitation of an injection molding grade Rilsan PA11 (Arkema) from ethanol at 120 °C and 20-wt.% polymer concentration. The process for particle production is outlined in great detail in [18]. The as-obtained powder after precipitation is characterized by a volume-averaged particle size of x50,3 = 90 μm and a x10,3 of 27 μm and a x90,3 of 185 μm. The particle size distribution (PSD) is rather broad, as expressed by the span (x90,3 - x10,3) / x50,3 of the PSD of 1.76. This material behaves cohesive (2 < ffc < 4). For improvement of flowability, the PSD was narrowed by sieving and the sieved PA11 powder with x10,3 = 34 μm, x50,3 = 84 μm, x90,3 = 165 μm (span = 1.56) was dry coated with 0.5 wt.-% of hydrophobic fumed silica, c.f.

red filled squares in Figure 3: the classified and functionalized powder now is easy flowing (ffc >4, HR = 1.21). Dry coating of micron-sized host particles with flowing aids (nano-sized guest particles) is a well-known approach to increase powder flowability [15, 17, 18, 23]. Briefly, the nanoscale guest particles act as spacers between the micron-sized host particles and, thus, reduce particle-particle interactions by lowering the van der Waals adhesion forces [14, 24, 25] being highly dependent on interparticulate distance.While shear testers allow for a well-reproducible determination of flowability, there is some criticism in using the obtained flow functions ffc for assessment of PBF processability, as ffc in shear testers typically is determined at rather high consolidation stresses that do not apply during powder spreading.

Moreover, the782(3)flowability of a pre-consolidated bulk solid is determined, which does not resemble the situation prior to powder spread in SLS machines. For example, if a rake is used for powder deposition, typically a non- consolidated bulk solid is spread; also if a roller coated is used, the bulk solid is applied as non-consolidated powder and the consolidation stress is rather low.classification non-flowing very cohesive cohesive easy-flowing free-flowingFlow function ffc / - < 1 1 … 2 2 … 4 4 … 10 > 10Hausner ratio HR / -> 1.4 1.25 … 1.4 1.1 … 1.25 < 1.1Table 1: Assignment of powder flowability: flow function ffc vs. Hausner ratio (HR) [21, 26].Powder tensile testerOne approach to determine flowability of non-consolidated powders is the determination of the powder tensile strength. The measurement principle is outlined in the sketch given in Figure 4, left and described in our previous work, e.g. [10, 23].

The measurement principle is outlined in the sketch given in Figure 4, left and described in our previous work, e.g. [10, 23]. A photograph of the powder tensile tester is depicted in Figure 4, right. The main components of the powder tensile tester are a powder reservoir, which is mounted on a motorized stage allowing for movement in vertical direction, a force transducer connected to a stamp that is moved into the powder reservoir and soft- and hardware that allows for force measurement data acquisition. First, the powder to be measured is prepared as non-consolidated bulk solid, which can be achieved by sieving the powder into the reservoir and the powder surface is smoothed by a rake. Then, the powder reservoir is moved towards the stamp being mounted on a traverse connected to the force transducer and covered with a thin layer of adhesive (to allow for adhesion of the powder particles in the uppermost layer), see Figure 4, left.

Upon contact with the powder layer, the stamp (slightly) penetrates the bulk solid until a certain (pre-set) maximal force is reached. Then, the motorized stage is stopped, the movement direction of the stage is reversed and the maximum force Fmax is registered. Due to the adhesive layer on the stamp of surface area Astamp, the uppermost particle layer is ‘glued’ to the stamp surface, thus, when moving back the stamp, the bulk solid breaks between the uppermost and the neighboring powder layer. The layer adhesion force Flayer-layer can be determined from the maximum force Fmax corrected for the gravitational force Fgravity due to the particle mass adhering to the stamp.

low powder tensile strength. An example for the application of powder tensile strength measurements for characterization of powders for PBF of polymers is given in Figure 5. There, tensile strengths determined for different polystyrene (PS) microparticles of comparable particle size distribution, namely comminuted, irregular shaped PS particles [27], thermally rounded PS particles and comminuted and rounded PS particles that have been subjected to dry particle coating with a flowing aid (fumed silica) [10] are given.

Right: photograph of the powder tensile strength tester.30 < ~ :o,.o25v0~6)/)t"' a. -01 ... -Ill -20~C: QI 15.!! Ill 10 C: QIOJ",?,Oti,<$)6 .1/ij1,• •••50edgedroundedrounded + 1wt% R1 + 1wt% R1edgedFigure 5: Effect of particle shape and surface functionalization on powder flowability as determined by powder tensile strength measurements: rounding of comminuted polystyrene (PS) particles and application of flowing aids by dry coating allows for reduction of interparticulate forces (reduction of powder tensile strength) and, thus, improvement of flowability [10].Dynamic tests: revolution powder analyzer, powder rheometer, powder spreading experimentsFor determination of flowability, different approaches have been discussed so far. For example, a so- called revolution powder analyzer has been proposed.

For example, a so- called revolution powder analyzer has been proposed. The device consists of a rotating cylinder being equipped with transparent (circular) walls that allow for optical determination of the avalanche angle of the powder in dependency of the rotational frequency of the cylinder. Thus, this method allows for assessment of flow under stress conditions that are similar to that during powder spreading, if (circumferential) rotational speed is set close to the translational speed of the powder spreading unit [1, 3, 22, 29]. Moreover, for characterization of flowability under dynamic conditions, the measurement of the fluidization behavior of the PBF materials (c.f. minimum fluidization velocity), funnel flow experiments or so-called powder rheometers [30] have been784proposed. In the powder rheometer, the resistance of a bulk solid against flow is assessed via measurement of the acting torques during movement of a rotating blade fixed to a shaft.

In the powder rheometer, the resistance of a bulk solid against flow is assessed via measurement of the acting torques during movement of a rotating blade fixed to a shaft. Depending on the mode of operation, the method allows for determination of cohesiveness and flowability, thus gives results that often correlate with HR or results from shear experiments. An extensive discussion and comparison of several methods can be found in the literature, e.g. [1, 3,22].Different powder spreader model setups have been proposed so far. For example van der Eynde et al. [28] proposed a device that consists of a powder spreading unit (rake with variable, known gap size) and a load cell connected to a measurement plate (of known area), which –if the solid density of the material is known- allows for determination of the bulk density of the spread powder in dependence of e.g. gap height or spreader speed.

gap height or spreader speed. Thus, this device allows for determination of a modified HR under powder spreading conditions (at ambient temperature). Powder spreading experiments on (typically black) model substrates with subsequent image analysis of the area covered with PBF material have been proposed to compare the spreadability of differently treated powders. For example, Blümel et al. [16] could show a good correlation between ffc determined by shear experiments, powder tensile strength, surface coverage with PBF material in the model powder spreading experiment and quality of laser sintered single layers for polyethylene powder coated with different flowing aids. Schmidt et al.

Schmidt et al. [17] could prove an analogous behavior for polybutylene terephthalate (PBT) powders: powder tensile strength, surface coverage with PBT powder in spreading model experiments and quality of SLS single layer specimen show a strong correlation as depicted in Figure 6.(2.9 +/- 0.6) PaFigure 6: Effect of particle shape (ground (left) vs. rounded and dry coated (right) of PBT powders on powder tensile strength (insets, top), surface coverage in the model powder spreading experiment ((19.1 +/- 4.2) % vs. (97.3 +/- 1.8) %, comminuted vs. rounded and dry coated) and quality of SLS single layers (bottom) [17].Tribo charging of PBF powders during powder spreadingA variant of the aforementioned powder spreading model setup has been recently proposed to assess the tribo charging of PA12 SLS powders of different ageing state during powder deposition.

The model deposition experiments were performed at room temperature, the setup is depicted in Figure 7, top; the spreader unit has been complemented by a probe of an electrostatic voltmeter, that allows for measurement of surface potentials and, thus, surface charge. The electrostatic potential was monitored in dependence on position, the number of consecutive application steps and the ageing state of the powder. Figure 7, bottom depicts exemplarily results for an aged PA12 powder: tribo charging during powder deposition occurs and a charge accumulation, i.e. increase of potential and, thus, charge with increasing number of powder applications was noted. The charge built up also is a function of ageing state of the SLS material (virgin powder shows less tribo charging). A detailed discussion can be found in Hesse et al.

[31].785WEiti e " • • • 8Electrostatic voltmeterPowder layerDoctorPET foil\----- bladePowder application140 > 120 -- t) cu :;:; :;:; 10 0 CU C u5 2 o o 80 .!::; Cl.. t) Q) © u 60 w~number of applications - -1 - - 2 - - 3 - - 4 - - 5::::i 40Cl)2020406080 1 00 120 140 160Layer position in application direction / mmFigure 7: top: Model setup for determination of tribo charging during powder spreading; bottom: electrostatic surface potential during spreading of aged PA12 SLS powder [31].Powder flowability at elevated temperaturesWhile, as outlined in the previous sections, there have been proposed many setups and measurement protocols to characterize powder flowability of PBF materials under different stress conditions, most of these experiments were performed at ambient temperature, i.e. little attention has been paid on the effect of temperature on flowability.

little attention has been paid on the effect of temperature on flowability. The effect of temperature on powder flowability is far from being negligible, as powder spreading in SLS is performed at building chamber temperature, which e.g. in the case of PA12 typically is around 170 °C. Moreover, information on flowability at ambient conditions does not allow for an assessment of flowability at elevated temperature, as particle interactions, which determine powder flowability, can show complex and pronounced dependencies on temperature, as demonstrated in studies for organic matter (lactose, coffee, waxes) [32] or fly ashes [33, 34]. For example, adhesion forces can increase with increasing temperature due to changes in adsorption layers, liquid bridges, contact areas due to visco-plastic deformation, or flowability can decrease due to formation of solid bridges with beginning of sintering [32-38].

In a recent attempt, a (temperature-controlled) rheometer was modified with an appropriate measurement geometry that allows to perform shear experiments at elevated temperatures similar to the protocols employed in shear testers (see above). This is exemplified in Figure 8, which depicts the angle of internal friction ϕϕsf as786determined in dependency on temperature for virgin PA12 powder (PA2200, EOS). ϕϕ sf is calculated from the shear stress (cid:2028) and the normal stress via ϕ sf = (cid:1853)(cid:1870)(cid:1855)tan ((cid:2028)/(cid:2026)). While steady state flow conditions are reached promptly for lower temperatures, at elevated temperatures (cid:2028) does reach a maximum value and starts declining rapidly afterwards. Therefore the evaluated states for higher temperatures are at maximum shear stress as opposed to steady state conditions for lower temperatures.

Obviously, a pronounced decrease of flowability is notable starting from around 150 °C up to a temperature close to the melting temperature of the polymer under consideration, as reflected by the increase by the increase in the angle of internal friction.65 - - r - - - - - - - - - - - - - - - - ,0virign PA12 (PA2200)060 t5 9- 55 ·;:: ~ ._ ~ 50 ro E ~ 45 _$ C --0 .Q 40 C ro Q) en 35 - 0) g' C 30 ro ·c :::J -o 25C 0'lii000000Q050100150200temperature / °CFigure 8: Temperature dependence of the angle of internal friction as determined for steady-flow conditions up to 150 °C and maximum shear stress above 150 °C for a PA12 SLS powder (PA2200, EOS).ConclusionsIn this contribution, different methods and experimental setups for assessment of bulk solid properties (flowability, packing density) relevant for processability of PBF materials were introduced and discussed. While the Hausner ratio (HR), i.e.

While the Hausner ratio (HR), i.e. the quotient of tapped density and (loose) bulk density, is frequently applied, one should keep in mind that the obtained results might be misleading, because HR is not a measure of powder flow, although for many materials a correlation of (small) HR and good flowability was found empirically. Moreover, the mode of powder consolidation and the measurement geometries can influence the result. Shear testers allow for reliable measurement of flowability (via the flow function ffc) under well-defined conditions; a positive correlation of large ffc value, small powder tensile strength and small HR was found for precipitated PA11 powder. Although, the stress conditions typically applied during the shear experiments are quite different from that during powder application. Consequently, powder spreading model experiments mimicking stress conditions during powder spreading are an alternative approach to characterize e.g.

Consequently, powder spreading model experiments mimicking stress conditions during powder spreading are an alternative approach to characterize e.g. surface coverage with SLS powders or the tribo charging of the materials. The aforementioned characterization methods, so far, typically were performed room temperature. Characterization of flowability at process temperature, which can be quite different from that at ambient conditions, remains a challenge for future work.AcknowledgementsThis study has been funded by the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation) – Project-ID 61375930 – SFB 814, Subprojects A1, A2 and T1). The financial support is gratefully acknowledged.787Literature [1] A. Amado, M. Schmid, G. Levy, K. Wegener, Advances in SLS Powder Characterization, SFFProceedings 2011, 438-452.[2] M. Schmid, A. Amado, K. Wegener, Journal of Materials Research 29 (2014) 1824-1832.. [3] A. B. Spierings, M. Voegtlin, T. Bauer, and K.

aracterization of Polyolefin – Alumina compounded mix for FDC processingV. V. Bhat1, K. Geetha1, R. N. Das3, B. Gurumoorthy2, A. M. Umarji1 1 Materials Research Centre, 2 Mechanical Engineering, Indian Institute of Science, Bangalore-560012, India. 3 Ceramic Technological Institute BHEL, Bangalore-560012, India.AbstractFused deposition of ceramics (FDC) uses thermoplastic binder and ceramic blend as feed material. The geometry of the fused deposition machine restricts the workability of the FDC feed material, which is in the form of a filament. The feasibility of the usage of these filaments is mainly based on the viscosity at the working temperature and the compressive modulus of the feed material at near room temperature. The polymers based on Polyethylene (PE), having two molecular weights 3,000 (LDWAX) and 341,000 (LDPE) were mixed in various weight proportions by solvent method using toluene as solvent, to develop a binder system for Fused Deposition Modeling of alumina.

Variation of viscosity as a function of composition, temperature and solid loading was measured using spindle viscometer and capillary rheometer. Dilatometric thermal expansion of 50 Vol% alumina compounded binders is measured up to 120(cid:176) C. The pronounced softening of the compounded mixture is observed beyond 70(cid:176) C, when the volume % of LDWAX is in excess of 50% total binder content. The compression strength decreases from 720 N to 310 N for pellet having 1.2cm diameter and 1.5 aspect ratio, when the percentage of LDWAX varies from 40 to 70% in the binder composition. The suitability of the compounded mix of LDWAX and LDPE binder with 40 Vol. % alumina for FDC is being evaluated in a StratasysTM1600 machine.IntroductionThe success of the FDC [1-2] or injection molding [3-4] is mainly dependent on the detailed characterization and selection of the proper feed material composition.

A multicomponent binder (MB) system containing base binder, plasticiser, and tackifier have been found to be useful in FDC [5]. The advantage of MB over simple unicomponent binder system is its cumulative and variable property with the composition. The different decomposition temperature of each component in the binder system will minimize the cracks in the sample after binder burn out. The role of materials processing [6] and solid loading are the other parameters, which control the property of the feed material. In this paper we restrict ourselves to a particular solid loading of 50 Vol. %.Though binder systems for FDC is being formulated for Al2O3, PZT, SiC etc.[7-8], there were very less instances of minimizing the number of efforts needed to formulate them. The viscosity at working temperature and compressive modulus of the feedstock are the main criterion in deciding the workability of a filament. This can be visualized in Fig 1.

The viscosity at working temperature and compressive modulus of the feedstock are the main criterion in deciding the workability of a filament. This can be visualized in Fig 1. The forward force exerted by the rollers and backward drag exerted by the viscous material in side the liquefier will lead to the bending of the filament in between the liquefier and roller [9]. To find out usefulness of the feedstock for FDC, the viscosity and compressive modulus measurements were carried out. The compressive modulus of the feedstock is being compared with the working investment casting wax supplied by Stratasys IncTM.42Forward force by rollerBackward drag due to viscosityFilamentLiquefierRollerResultant (buckling of filament)NozzleFig 1. Schematic Structure of roller liquefier array in twin head Stratasys IncTM. FDM-1600 machine.PE of two different molecular weights has been blended by solvent method and their viscosity and compression strength values are reported here.

FDM-1600 machine.PE of two different molecular weights has been blended by solvent method and their viscosity and compression strength values are reported here. The other parameter, which adds to the failure of the filament, is the softening of the filament at inlet of the liquefier. The dilatometric study is done to measure the softening of the material.ExperimentalThe molecular weight determination of Low Density Polyethylene (LDPE) and LDPE Wax (LDWAX) was determined out using HPLC-GPC system (Water Inc.). The binder system with LDPE as main binder and LDWAX as plasticiser was blended in different volume (Table 1.) fractions using toluene as the solvent. The blending was performed at 90(cid:176) C using a mechanical stirrer until the solvent almost removed. The polymer mass was transferred to a tray of aluminum foil. The solvent traces were removed by degassing at 110(cid:176) C and -650 mmHg in a vacuum oven.

The polymer mass was transferred to a tray of aluminum foil. The solvent traces were removed by degassing at 110(cid:176) C and -650 mmHg in a vacuum oven. The degassed sample was subjected for viscosity measurement using Brookfield viscometer (model- DV-II) with spindle #34, at varied temperature and revolution per minute.Table 1. The code for binder and ceramic binder blend with varying plasticiser content.Plasticiser content in the total binder (Vol.%) 0 40 50 60 70 80 100Identity Without Ceramic With 50 Vol. % Alumina Loading ----- CLDWAX 40% CLDWAX 50% CLDWAX 60% CLDWAX 70% ------ ------LDPE LDWAX 40% LDWAX 50% LDWAX 60% LDWAX 70% LDWAX 80% LDWAXThe 50 Vol.% solid loaded thermoplastic binder containing varying ratio of LDPE and LDWAX with alumina were prepared using solvent method. The average mean particle43size of the alumina was found to be ~0.6m m, (specific surface area: 8-10 sq.m/gm).

The average mean particle43size of the alumina was found to be ~0.6m m, (specific surface area: 8-10 sq.m/gm). Alumina powder was ball milled for one hour with 2 Wt.% of strearic acid in toluene medium. The surfactant treated alumina was completely transferred to a stainless steel container and added with calculated amount of LDPE and LDWAX. All the additives were mixed at 100(cid:176) C using mechanical stirrer. After almost all solvent has been removed, the mixture was transferred to the tray of aluminum foil and degassed at 110(cid:176) C and –650 mmHg using vacuum oven. The mixture was granulated and its rheology was measured using capillary rheometer CFT-500C type (Shimadzu Corporation) at different temperature with different shear rate.The pellets having 12mm diameter and 15mm height were made out of binder compounded with alumina at the load of 50kN in 100(cid:176) C.

Investment casting wax (ICW-05, Stratasys Inc TM.) pellets were made into the same dimension pellets as alumina by melting it in the same dye. The crushing strength of these pellets were measured in Darteck-9500 mechanical analyzer with the plunger speed of 0.001mm/s. The thermal expansion measurement of these pellets were carried out in home made dilatometer up to 150(cid:176) C.The filaments of 1.7mm of the alumina-binder mix were made using capillary rheometer at 100(cid:176) C. The extruded 1.5m length filaments were stored in a stainless steel tube till actual fused depositing on the Stratasys 1600 machine. The liquefier fitted with stainless steel nozzles was maintained at 130(cid:176) C and the filament was introduced into the machine with the flow rate of 40% (Stratasys specification).Results and DiscussionThe decrease in the chain length of the polymer is directly proportional to the interaction between two polymer chains.

Therefore the viscosity is intern related to the average molecular weight of the binder. Table 2. shows the decrease in the average molecular weight of the binders and their viscosity at 150(cid:176) C. This is an important data in controlling the viscosity of the binder on the basis of average molecular weight. We could measure the viscosity of the polymer composite in the range, which is shown in Fig 2, rest of the valves was out of viscometer range. These sets of thermoplastic binders have shown clear expected pseudoplastic behavior in the whole range of temperature, which indicates that there is no effect on the nature of the binder by solvent mixing.Table 2.

The average molecular weights of binder plasticiser mixture at different proportion.Polymer Composite (% of LDWAX)Molecular Weight Average (Mw)Viscosity at 150(cid:176) C (cps)0 40 50 60 70 80 100341000 205800 172000 138200 104400 70600 3000HIGH 85300 33900 14400 5510 1800 LOW44The viscosity of surfactant treated alumina with binder having variable volume fractions of binder and plasticiser showed a characteristic pseudoplastic behavior as expected. The variation of viscosity as a function of shear rate at different temperature is shown in Fig 3. All the ceramic-binder blends with different plasticiser content were found to have viscosity less than 8000cps at particular temperatures corresponding to them. The preferred working temperature is below 150(cid:176) C, to avoid the plasticiser loss due to evaporation.Fig 2. Variation of viscosity at different temperature for different average molecular weightbindersFig 3.

Variation of viscosity at different temperature for different average molecular weightbindersFig 3. The viscosity behavior of binder with alumina as a function of shear rate at differenttemperature.45The increase in the plasticiser above 50 Vol.% of binder drastically brings down the long chain interaction in the polymer which causes dramatic change in the viscosity of the material. This is clearly evident from the Table No: 3. There is minimum change in the viscosity of CLDWAX-50% and CLDWAX-40%.Table 3.

This is clearly evident from the Table No: 3. There is minimum change in the viscosity of CLDWAX-50% and CLDWAX-40%.Table 3. Variation in viscosity at 130(cid:176) C (at different percentage of plasticiser in the totalbinder content.) as a function of plasticiser content in the total binder system.Shear Rate1333.33 666.67 266.67 133.33 66.67CLDWAX-60% 591 866 1506 2516 4447Viscosity (cps) CLDWAX-50% CLDWAX-40%1306 1931 3350 5325 91871691 2440 3950 6012 9187The increase in the plasticiser content can bring down the viscosity of the material but at the same time decreases the mechanical property, which adversely affects the workability of the filament. This is because of sagging of the filament near the inlet of the liquefier and low compressive modulus of the material.The thermal expansion measurements with the constant load of 55g shows, at low temperature stiffness of the materials decreases with increasing plasticiser content.

This is evident from the Fig 4., the sagging of the sample starts in the range of 65-70(cid:176) C for the binder having LDWAX more than 60%. But the binder having 50% of LDWAX show the sagging after 80(cid:176) C. These type of behavior is due to the short range glass transition of the plasticiser at around 55(cid:176) C. In the case of binders having 40% plasticiser, the sagging will not be found till it softens at100(cid:176) C. All the samples will collapse at 100(cid:176) C due to softening of LDPE, during glass transition.Fig 4. The thermal expansion behavior of 50 Vol% of alumina solid loaded binder withdifferent proportion of plasticiser.The binder ceramic mixtures were found to have the modulus ranging from 1.1kN/mm to 0.4kN/mm. The working filament ICW-05 filament supplied by Stratasys46IncTM. was found. to has the modulus of ~0.55 kN/mm as shown in Fig 5.

The working filament ICW-05 filament supplied by Stratasys46IncTM. was found. to has the modulus of ~0.55 kN/mm as shown in Fig 5. The ceramic binder mixture which are having plasticiser content less than 50% have higher strength and modulus than ICW-05. Though CLDWAX 40% and 50% show higher modulus and compressive strength, they are brittle compared to the ICW-05. The ceramic binder blends show sudden softening once they attain the maximum elastic modulus. The modulus and compressive strength of the ceramic-binder blend and ICW-05 are presented in Table 4. But the ICW-05 shows the deformation in the shape after its maximum elastic modulus, while the sample will be compressed with increasing load. Ceramic-binder can withstand more load compared to ICW-05 without compression, whereas ICW-05 can withstand higher load even after it starts deforming. This tough behavior of the ICW-05 is advantageous over the ceramic-binder blend.

This tough behavior of the ICW-05 is advantageous over the ceramic-binder blend. The use of other additives such as tackfier, microcrystalline wax to the binder system may help in overcoming this problem.Table 4. The comparison in between the modulus and compressive strength of ceramic-binderblends with ICW-05.Materials compositionICW-05 CLDWAX 40% CLDWAX 50% CLDWAX 60% CLDWAX 70%Modulus(N/mm) 550 1100 820 475 410Compressive Strength {until deformation} (N) 570 720 610 350 310Fig 5. The comparison of the modulus of various ceramic binder (50% solid loading) mixturewith ICW-05.The viscosity of ceramic-binder blend can be adjusted to the required level by controlling the temperature. The mechanical analysis show that the compressive modulus and47the strength of CLDWAX 60% and CLDWAX 70% are lower than ICW-05.

The mechanical analysis show that the compressive modulus and47the strength of CLDWAX 60% and CLDWAX 70% are lower than ICW-05. The dilatometric curves have shown that the sagging of CLDWAX 60% and CLDWAX 70% will start at 65- 70(cid:176) C, which is not advisable because of their sagging near the inlet of the liquefier. On the basis of these results CLDWAX 40% and CLDWAX 50% have been chosen for the testing on Stratasys machine. The liquefier was maintained at 140(cid:176) C and the filament was introduced in the speed of 40% (Stratasys Inc. calibration). The filaments of CLDWAX 40% could be extruded through the machine without buckling.ConclusionThe polyethylene based multicomponent binder system was formulated for the 50 Vol% solid loading. The ceramic-binder blends were characterized by viscosity, dilatometric thermal expansion and mechanical strength measurements to narrow down on a particular composition.

The ceramic-binder blends were characterized by viscosity, dilatometric thermal expansion and mechanical strength measurements to narrow down on a particular composition. The results have shown CLDWAX 40% and CLDWAX 50% are having the required viscosity at 140 and 130(cid:176) C, and having better mechanical properties than ICW-05. The filaments of CLDWAX 40% could be extruded through the liquefier of FDM. Increasing the number of components of the multicomponent binder system to include tackifier, microcrystalline waxes are being attempted to evolve composition to make FDC process smooth.Acknowledgements We acknowledge Mr. P. Joshi for helping us in performing viscosity measurements in BHEL. We sincerely thank Mr. Shashidhara of Dept. of Metallurgy, IISc., for his timely help in conducting compression strength measurements. We also acknowledge Mr. Girish’s help while handling FDM machine.Reference1. Agarwala, M. K., Bandyopadhyay, A., Weeren, R., Safari, A., Danforth, S.

C., Safari, A. “Deposition of Lead Zirconate Titanate for Fused Deposition of Ceramics” Private Communication.9. Venkatraman, N., Rangarajan, S., Mattewson, M. J., Safari, A. and Danforth, S., C. “Mechanical and Rheological Properties of Feedstock Material for Fused Deposition of Ceramics and Metals (FDC and FDMet) and their Relationship to Process Performance” SFF proceedings-1999, p.35149.

ARACTERIZATION OF SELECTIVE LASER SINTERING MATERIALS TO DETERMINE PROCESS STABILITYT.J. Gornet, K.R. Davis, Dr. T.L. Starr, K.M. MulloyRapid Prototyping Center, J.B. Speed Engineering School, University of Louisville, Vogt Building, Room 101, Louisville, KY 40292AbstractThe Selective Laser Sintering (SLS) process has proved to be an excellent method for prototyping functional parts out of engineering thermoplastics such as polyamides. However, the material undergoes physical and chemical changes due to repeated heating cycles in the SLS equipment. This causes variations in powder characteristics and performance in the SLS process. With the increased utilization of SLS for direct manufacturing it is necessary to develop a characterization and testing system that can determine powder fitness to ensure process stability and part quality. Current powder recycling methodologies use an average virgin-to-used powder mixture.

Current powder recycling methodologies use an average virgin-to-used powder mixture. In a new approach, a testing mechanism to deliver a numerical, measurable material characterization will be discussed. Experimental results of repeated reuse of material and its resulting physical effects on mechanical properties, shrinkage, and chemical tests will be presented. A definitive testing and measurement process control will be shown to improve process stability and thus part quality and consistency.IntroductionThe 3D Systems Corporation, Valencia, CA, manufactures and sells Selective Laser Sintering (SLS) equipment and supplies powdered materials for use in this equipment. DuraFormTM is a polyamide polymer with excellent mechanical properties and is used to produce functional engineering models and prototypes.

DuraFormTM is a polyamide polymer with excellent mechanical properties and is used to produce functional engineering models and prototypes. 3D Systems purchases this powder from a third party and supplies it to customers.3D Systems and its customers have observed two different problems with DuraForm powder since its introduction: batch-to-batch variation in virgin material and progressive variation of powder performance after repeated use and reblending. As the SLS process transitions from functional prototyping and is incorporated into direct manufacturing processes, part quality and consistency is becoming a more important factor.In an ideal process, all of the used powder could be recycled to achieve 100 per cent utilization. However, just as in injection molding, the repeated heat cycles change material characteristics resulting in part quality issues. Starting with virgin powder, the builds are very good with excellent surface finish and feature definition.

Starting with virgin powder, the builds are very good with excellent surface finish and feature definition. As the powder is reused, the surface finish degrades and at some point becomes unacceptable. Therefore, virgin powder is blended into the mix in an attempt to keep the part quality at an acceptable level. SLS system users have generally blended their powders in the following manner:1. Run a build.5462. Break out the parts, segregating the unsintered powder that can be easily broken up as “used” powder.3. Sift the used powder in a vibratory sifter4. Blend in a weight percentage of virgin powder and mix via mechanical methods.The weight percentage varies from user to user and material to material. For DuraForm users generally blend 25 to 33 percent virgin powder. Powder management and blending is a critical part of the SLS process due to the high cost of SLS powders on a per pound basis.

For DuraForm users generally blend 25 to 33 percent virgin powder. Powder management and blending is a critical part of the SLS process due to the high cost of SLS powders on a per pound basis. Blending too much virgin powder into the process raises costs and blending too little increases the chance of unacceptable surface finish and part detail.ObjectiveThe University of Louisville Rapid Prototyping Center proposed to investigate physical and chemical characteristics and mechanical property changes of DuraForm powder. The goals of this investigation were to identify those characteristics that correlate with poor performance and to develop an evaluation protocol that will ensure consistent, good performance.

The goals of this investigation were to identify those characteristics that correlate with poor performance and to develop an evaluation protocol that will ensure consistent, good performance. From those results it was desired to develop a cost effective, efficient measurement and testing procedure for blending that will provide a metric that can aid in the process control of SLS as a production process.MethodologyRepresentative PartsA build packet for a 3D Systems SLS2000 was created that utilizes the full build volume and the entirety of the powder supply pistons. A number of different parts were chosen to highlight surface finish in all three axes, fine feature detail, scale and offset parameters, and tensile specimens. A total of 48 parts were in the build. The entire build packet is depicted in Figure 1. The overall build height including warm-up was 11.9 inches.Figure 1.

A total of 48 parts were in the build. The entire build packet is depicted in Figure 1. The overall build height including warm-up was 11.9 inches.Figure 1. SLS2000 Build PacketProcessing and Powder AcquisitionThe initial step in the process was to determine the best operating parameters for the virgin powder: laser power, feed temperatures, part bed temperature, and piston heater temperature. These critical parameters would remain constant for all of the future builds to547provide consistent heat input to the powder. These parameters can vary from machine to machine. Three containers of DuraForm from the same lot were selected for the trial. A sample of the virgin powder was taken and labeled. The SLS2000 feed cartridges were loaded with powder. The feed rate was such that the build packet would use the entire supply of powder, thereby exposing it all to a single heat history. The complete build was run and all parts were broken out.

The feed rate was such that the build packet would use the entire supply of powder, thereby exposing it all to a single heat history. The complete build was run and all parts were broken out. All of the unsintered powder was broken up and sifted through a vibratory sifter. The overflow cartridge powder was added to the part cake powder and mechanically mixed thoroughly. A sample of this mix was taken and labeled.The resultant remaining powder was put back through the SLS process using the same build packet. The volume of powder used by the parts decreased the amount of powder available for the second build. Therefore, the build would not complete the full height. The build decrease in height was approximately 12.5 percent. The build was run until the build terminated due to lack of powder. Again, the parts were removed from the part cake and labeled.

The build decrease in height was approximately 12.5 percent. The build was run until the build terminated due to lack of powder. Again, the parts were removed from the part cake and labeled. All powder was sifted, mixed, and a sample was taken and labeled.This process was repeated until the build conditions deteriorated or there was too little powder remaining. This was a total of seven consecutive builds for DuraForm.Physical and Chemical TestingSeveral tests were planned for the part samples and powder specimens. The interest wasin examining the possible correlation between part quality and several test criteria to indicate material stability and degradation in the process.Melt-Index determination: Melt-index was measured using an extrusion plastometer according to ASTM D1238. This index is a measure of the flow characteristic of the molten polymer and is sensitive to differences in the basic polymer structure due to changes in molecular weight.

This index is a measure of the flow characteristic of the molten polymer and is sensitive to differences in the basic polymer structure due to changes in molecular weight. Changes in melt-index should correlate with changes in the build characteristics of DuraForm during the laser sintering process.Differential scanning calorimetry: Differential Scanning Calorimetry (DSC) was used to compare the melting and solidification behavior of the powders. Detailed run parameters were established during preliminary testing and included a 100% nitrogen environment, heating at constant rate to a set temperature above the melting point of the material and cooling at a fixed rate to room temperature.

The resulting thermal scans are compared in order to identify differences in features such as melting or solidification onset temperature, melting or solidification range and heat of fusion.Mechanical properties: Mechanical properties of test bars of the materials were measured using an Instron test machine. Measurements included tensile yield strength, rupture strength and elongation. Five standard ASTM tensile specimens were built with the long axis oriented along the X direction and five along the Y direction.Surface appearance and detail: A subjective evaluation was used to compare similar parts on a run-by-run basis. The evaluation was used to determine the lowest acceptable part quality for a user. This result gave the baseline for end-use testing.548ResultsMelt IndexThe melt index testing was conducted using ASTM standard procedures for nylon materials. A temperature of 235°C and a weight of 2.16 kg were used for the DuraForm material.

A temperature of 235°C and a weight of 2.16 kg were used for the DuraForm material. The results are shown in Figure 2. It can be seen from the graph that the viscosity of the material decreased with repeated heat exposures from the SLS process. The melt index test is a good indicator of changes in molecular weight. In this test, the results indicate an increase in the molecular weight of the material with multiple heating cycles. In injection molding of thermoplastics, the melt index will generally increase with multiple thermal exposures indicating a decrease in molecular weight.Melt Index Change over 7 Builds50) n m 0 1 g (i/4030R F M20100Ivirgin1234567BuildMFRFigure 2. Melt Index ResultsDifferential Scanning Calorimetery (DSC)Approximately 10mg of the sample powder is deposited into the chamber of the DSC equipment and the chamber is purged with nitrogen gas.

Melt Index ResultsDifferential Scanning Calorimetery (DSC)Approximately 10mg of the sample powder is deposited into the chamber of the DSC equipment and the chamber is purged with nitrogen gas. The temperature is equalized at 75°C and then ramped at five degrees per minute to 225°C. Figure 3 shows a representative DSC output curve illustrating the melt point and the recrystallization points. The melt points for all of the runs are shown in Figure 4. The trend of the curve indicates that the melt point is increasing with multiple heat exposures.549Figure 3. DuraForm DSC CurveMelt Point Change over 7 Builds188s u s e C g e dil.187.5187186.5p m e T186185.5Ivirgin1234567BuildMelt Pt.Figure 4. Melting PointMechanical PropertiesTensile strength and elongation results are shown in Figure 5. The tensile strength increases slightly over the first five builds.

Melting PointMechanical PropertiesTensile strength and elongation results are shown in Figure 5. The tensile strength increases slightly over the first five builds. However, there is an approximately 25 per cent drop off in tensile strength at the sixth build. The tensile bars aligned along the Y-axis exhibited slightly greater tensile strength.550Tensile Strength and % Elongation) IS P( h t g n e tS e l i s n e T6000500040003000200010009 8 7 6 5 4 3 2 1 0)%( n o i t a g n o Eltensile x tensile y % elongation x % elongation y01234567BuildFigure 5. Tensile Strength and ElongationSurface FinishThe evaluation of the surface finish was performed as a qualitative assessment. A surface finish commonly called “orange peel” is a known phenomenon that is associated with over used or degraded powder. The determination of an acceptable surface finish varies from user to user. Figure 6 shows parts with good and bad surface finish.

The determination of an acceptable surface finish varies from user to user. Figure 6 shows parts with good and bad surface finish. The right view is a classical “orange peel” finish. Visual evaluation of parts from the seven builds indicate that the surface finish is excellent for the first three builds, acceptable for the next two builds, and unacceptable after that.Figure 6 Part Surface Finish ComparisonScale and OffsetScale and offset test coupons were run in each build to track dimensional changes that can occur as the powder ages. This is an important test to understand if dimensional inaccuracies551are occurring due to variations in powder quality. The parts were measured and the results input into a scale and offset worksheet supplied by 3D Systems to track these values. Figure 7 is a graph showing the changes over the seven runs. The scale in the X direction varies from 2.8% to 3.4% with the Y scale varying from 3.0% to 3.5%.

Figure 7 is a graph showing the changes over the seven runs. The scale in the X direction varies from 2.8% to 3.4% with the Y scale varying from 3.0% to 3.5%. The offset ranges from 0.0057 to 0.0022 inches and from 0.0037 to 0.0005 inches in the X and Y directions, respectively.Duraform Study Scale and Offset for 7 buildse l a c S1.036 1.034 1.032 1.030 1.028 1.026 1.0240.007 0.006 0.005 0.004 0.003 0.002 0.001 0.000t e s f f ox-scaley-scale1234567x-offsetBuild #y-offsetFigure 7. Scale and Offset ValuesConclusionsThe most notable conclusion from these sets of results is that powder variation on a run- to-run basis can significantly impact the quality of the produced parts. While some variation may be acceptable for prototype parts, it is an unacceptable state for direct manufacturing.

While some variation may be acceptable for prototype parts, it is an unacceptable state for direct manufacturing. This fact alone illustrates the need for an effective powder management and testing system to control critical parameters.The combined indications of the melt index test and the differential scanning calorimeter suggest a change in the polymer chemistry of the DuraForm powder exposed to multiple heat cycles. The decrease in the melt index value points to an increase in the molecular weight. Combine that with the increase in the melt point from the DSC equipment and it appears as if there is a limited “cross-linking” of the material similar to what might be seen in some thermoset material systems. While the DuraForm is a known nylon-based thermoplastic, it appears as if there is some lengthening of the polymer chains. This is the opposite of what normally occurs in many thermoplastics during the injection molding process.

This is the opposite of what normally occurs in many thermoplastics during the injection molding process. Repeated heat cycles or over exposure due to barrel residence times generally causes a decrease in molecular weight and an increase in the melt index. The material properties suffer accordingly.The challenge was to find a cost effective method for tracking the fitness of the DuraForm powder. It is desired to have consistent powder properties in the direct manufacturing process to allow process temperature controls and scale and offset values to remain constant. The melt flow rate equipment has proved to be a good predictor and provide a metric by which to track the powder and check its condition.552The first requirement is to define a target melt index parameter at which to run the SLS process. This value can be extrapolated from the test builds by evaluating the surface finish of the parts produced during the runs.

This value can be extrapolated from the test builds by evaluating the surface finish of the parts produced during the runs. The physical strength of the part, tensile strength and elongation, does not fall off until after a point most users would consider the surface finish to be unacceptable. In addition, by using the melt index as the control, the user can control their surface finish at a specific level depending upon needs and part use.Tracking the melt index after each individual build has shown a range of degradation during the build. Builds that have smaller cross sectional scans per layer show a smaller change in the melt index compared to “hotter” builds that have longer scan times and thus longer exposure of the powder to the heaters. Therefore, the amount of virgin DuraForm to be added to achieve target melt index has varied from 15 to 40 percent.

Therefore, the amount of virgin DuraForm to be added to achieve target melt index has varied from 15 to 40 percent. Blending curves must be generated to create a method for determining the exact percentage of new material by weight to be mixed with the used material. Starting with used powder and adding varying amounts of virgin powder and then retesting for the melt index accomplished this. By populating a dataset with numerous blends it was possible to create a blending curve and resulting equations for a specific target melt index.Since the inception of this testing procedure, the University of Louisville Rapid Prototyping Center (RPC) has seen very consistent results. Part quality has been reproducible build-to-build and the scale and offset numbers have remained constant. It has been a worthwhile tool to for incoming inspection of material and for determining the quality of unknown lot of used DuraForm powder.

It has been a worthwhile tool to for incoming inspection of material and for determining the quality of unknown lot of used DuraForm powder. For example, if the melt index testing shows that more than 50% of virgin DuraForm is required to reach the target value, it can be discarded. Previously a visual, touch and feel method was used to find out which powder lots to discard. By tracking the mechanical properties and scale and offset numbers with each build, the SLS equipment may also be benchmarked. A small drift in the scale and offset was noticed, with no change in the melt index. Normally, the material would be blamed. Since the material quality was known, it was determined that it was a hardware issue with the equipment. This can assist in tracking the capability of the process and the machine.As the SLS process continues to transition into the new realm of direct manufacturing offunctional plastic parts, methods for controlling the process parameters are required.

For the DuraForm material, the melt index has proved to be a good indicator of material fitness and a control parameter for the process. Additional work is currently in progress to determine similar testing for other SLS materials such as glass-filled DuraForm.Selected ReferencesChoren, J, “SLS Powder Life Study”, 2001 Solid Freeform Fabrication Symposium, pp. 39-44.DeGuglielmo, Joseph, “New LNC-7000 Powder Recycling Procedure at Kodak”, presented at the 1997 SLS User’s Group, Albuquerque, NM, October 6, 1997.Leigh, David, “SLS Powder Management”, presented at the 2000 SLS User’s Group, New Orleans, LA, September 25, 2000.553.

ARACTERIZATION OF SHRINKAGE AND STRESSES IN LARGE AREA MASKLESS PHOTOPOLYMERIZATION Kiran Kambly, Dajun Yuan, Peng Shao, Suman Das* Woodruff School of Mechanical Engineering, Georgia Institute of Technology, Atlanta 30332ABSTRACT Large Area Maskless Photopolymerization (LAMP) is a high throughput direct digital manufacturing technology being developed for producing ceramic investment casting molds. Polymerization shrinkage and accompanying stresses developed during photopolymerization of ceramic particle-loaded resins in LAMP can cause deviations from the desired geometry. The extent of deviations depends on photoinitiator concentration, filler loading, degree of monomer conversion and operating parameters such as energy dose. An understanding of shrinkage and stresses built up in the part can assist in developing source geometry compensation algorithms and exposure strategies to alleviate these effects.

An understanding of shrinkage and stresses built up in the part can assist in developing source geometry compensation algorithms and exposure strategies to alleviate these effects. Real-time Fourier Transform Infrared Spectroscopy (RTFTIR) operated in Attenuated Total Reflectance (ATR) mode is used to characterize the three-dimensional shrinkage stresses. This work is sponsored by DARPA Grant HR0011-08-1-0075.INTRODUCTION Diacrylate based ceramic particle-loaded resins with appropriate photoinitiators form the basis for ultra fast photopolymerization reaction in LAMP. In the beginning of the reaction, the monomer molecules exist at Vander Waals distance which is of the order of ~104 A0. As the reaction proceeds, these monomer molecules form a closely packed network there by reducing the Van der Waals distances between them to covalent bond lengths (~ 1A0). This formation results in cumulative bulk contraction as the part is fabricated layer over layer by means of LAMP.

This formation results in cumulative bulk contraction as the part is fabricated layer over layer by means of LAMP. The degree of shrinkage is directly a measure of number of such covalent bonds formed. Thus shrinkage in LAMP can be characterized by measuring the number of covalent bonds formed during the reaction layer over layer. Some of the older techniques like Photo Differential calorimetry (Photo-DSC) made use of the exothermic nature of the reaction to determine the extent of reaction. Since the reaction time is extremely small of the order of 20~40ms in LAMP, it’s not feasible to pick up the heat of reaction within such small time scale. Several mechanical techniques likes Bonded Disk Measurement device and water dilatometers were used to characterize axial shrinkage during the reaction. However the in-plane shrinkage is assumed to be negligible which is not true for LAMP. Hence it is desired to have an analytical tool to monitor shrinkage layer over layer.

However the in-plane shrinkage is assumed to be negligible which is not true for LAMP. Hence it is desired to have an analytical tool to monitor shrinkage layer over layer. This data can further be used with the slicing algorithms in LAMP to accommodate for the shrinkage compensation.Fourier Transform Infrared Spectroscopy (FTIR) FTIR Spectroscopy is an analytical tool used to measure various chemical species and bonds in a molecule. It is one of the most powerful analytical methods for monitoring UV-initiated curing processes which proceed within a fraction of a second. It has several advantages over other methods such as Photo-DSC. The most important limitation of Photo-DSC is its long response time so that it is not possible to monitor polymerization reactions which occur within less than 10 s (Wu and Halloran 2005).

The most important limitation of Photo-DSC is its long response time so that it is not possible to monitor polymerization reactions which occur within less than 10 s (Wu and Halloran 2005). In addition, the Photo-DSC method requires knowledge on thetheoretical enthalpy of reaction to calculate the conversion of functional groups from the heat release measured. The sample thickness is much higher than that in most practical applications (≤500µm). Moreover, the thickness of the layer in the aluminum pan is poorly controlled. In contrast, FTIR spectroscopy allows a rapid and quantitative measurement of the conversion of specific reactive functional groups under conditions (light intensity, photoinitiator concentration, coat thickness, etc..) which are closely matched to those in technical coating and printing processes than it can be done with Photo-DSC.

In the past, FTIR spectroscopy has been successfully used to study the kinetics of photopolymerization reactions in dependence on the irradiation conditions and other experimental parameters, the reactivity of monomers and the efficiency of newly developed photoinitiator systems (Scherzer 2002). In the present study, the performance of photoinitiator systems under irradiation with monochromatic light and the kinetics of the proceeding ultrafast photopolymerization reaction were studied by FTIR spectroscopy to monitor conversion and polymerization shrinkage. Moreover, the effect of diverse physical and chemical factors on the kinetic behavior will be discussed and the extent of shrinkage will be estimated. Fig 1, represents FTIR operated in attenuated total reflectance mode, where in the sample is held by a pressure device in contact with a crystal, while the IR beam is made to strike the sample with a set of mirrors.

This results in a spectrum which is a unique to a specific composition and hence can be regarded as the molecular finger print of the sample as in no two unique molecular structures produce same spectra.Fig 1. FTIR operated in Attenuated Total Reflectance mode .During the polymerization of acrylate monomer, the carbon double conversion takes place by means of stretching and twisting (Wu and Halloran 2005), Stretching is sensitive at the 1640- 1610 cm-1 while twisting is sensitive at 1410 cm-1 as can be seen from the peaks in Fig 2. The conversion of double bonds during photopolymerization is captured as a function of time or the energy dose for the reaction. With increasing the reaction time, the peak heights are reduced and flatten out indicating the progress of the reaction. In order to quantify the amount by which the carbon double bond peak decreases, it’s necessary to compare it with a standard peak which do not change with respect to time.

In order to quantify the amount by which the carbon double bond peak decreases, it’s necessary to compare it with a standard peak which do not change with respect to time. In other words, it’s essential to have a non participating species in the composition. For the acrylate monomer, the carbon oxygen bond is taken as the reference peak since it remains constant with time. The degree of conversion α(t) is obtained using Eq (1).Fig2.

For the acrylate monomer, the carbon oxygen bond is taken as the reference peak since it remains constant with time. The degree of conversion α(t) is obtained using Eq (1).Fig2. FTIR spectrum of acrylate monomer undergoing Photopolymerization (Wu and Halloran 2005)α(t)=  Area     0 +Area+AreaArea+Area1620 Area14101620 Area   1750   1750Area1620 Area1410175001410  tX100(1)The rate of polymerization can be obtained by numerically differentiating Eq (1).Theoretical relationship between shrinkage and Degree of conversion In the polymer science literature it is widely accepted that the dominant cause of shrinkage-strain in acrylates arises from conversion of C=C double bonds, where for each monomer segment of the chain the larger van der Waals inter-molecular spacing is replaced by the smaller intra- molecular covalent bond (Hofmann, Denner et al. 2003; Watts 2005)).

Here Mm is the molecular weight and ρm density of the monomer. Thevalues of density and molecular weight were obtained from Material data sheet of ciba.In the more general case of multiacrylates, where f is the functionality of the monomer, thenumber of functional groups present in volume (V) is  f XVρ Mmm  .The number of functionalgroups reacted in volume (V) is α(t) x  f XVρ Mmm  .The percentage relative change in volume (volumetric shrinkage-strain) is∆V V()% = 22.5 X α(t)X  f XVρ Mmm  X 100(2)For a mixture of monomers of any functionality, ∑ ∑ (( ) f χ i i M χ mi∆V V()% = 22.5 X α(t)iρ X100)mixiiwhere fi is the functionality of monomer (i), χi is mole fraction of monomer (i), Mmi is molecular mass of monomer (i) and ρmix is density of the monomer mixture.

For example, with two monomers, one di-functional (Hexanediol diacrylate, HDDA) and another tetra functional (Ethoxylated penta erythryitol tetra acrylate, EPETA),(3)∆V V()% = 22.5 X α(t)2χ +4χ 1 M ρ1+M ρ22m1m2ρ X100mix(4)Eq 2-4 embody the expectation that the volumetric shrinkage-strain will be directly proportional to the degree of conversion of the monomer systems. Hence if the conversion diminishes for any reason, the shrinkage-strain should also be observed to decrease. Unless significant local porosity is induced within the material bulk, the local shrinkage-strain will translate into an observable macroscopic shrinkage-strain.Material system for LAMP Typical composition for LAMP consists of mixture of monomers with ceramic particles as filler content and a dispersant to prevent clogging of these ceramic particles.

Photoinitiator system is chosen such that it matches with the spectral characteristics of the UV source used for maximal utilization of the free radicals. Absorbers could further be used to alter the cure depths based on the process requirement. HDDA with EPETA were mixed in the ratio of 9:1. HDDA is a low volatile, low viscosity monomer, while EPETA is a fast curing monomer. Silica particles with 7µm mean diameter and 95% of them with less than 25µm were chosen as filler particles and variquat CC59 as dispersant. Irgacure 184 from sartomer was used to initiate the polymerization reaction.Experiments The objective of reducing the shrinkage can be achieved by systematically investigating effects of some of the key components of the material system. They are photoinitiator concentration, filler loading and the energy dose or the exposure time.

They are photoinitiator concentration, filler loading and the energy dose or the exposure time. Low exposure time results in poor layer to layer bonding while high exposure times lead to print through, over curing and side scattering effects. Very low filler content results in poor mechanical properties while very high filler content restricts the mobility of the monomer molecules and hence the conversion. Optimizingthese parameters is critical in order to achieve overall minimum processing time for LAMP. With these objectives, two set of experiments was designed. Firstly, the effect of photoinitiator concentration was studied by treating this as a variable and filler loading % was kept constant and vice versa in the second set. The details are listed in Table 1. Table 1. parameter details for the experiments.

The details are listed in Table 1. Table 1. parameter details for the experiments. Set PI concentrationExposure time (ms) 35 to 100Power input (W/cm2) 1.64Filler Loading (Vol%) 60(%) 0.2 to 0.6120.4360 to 851.6410 to 80Starting exposure times are so chosen such that it is just enough to begin photopolymerization. The samples thus prepared on the LAMP were used on FTIR spectrometer taking enough precautions such that they are prevented from undergoing further visible light curing. This enables to determine contraction purely due to photopolymerization on LAMP.Results and Discussion Effects of Photoinitiator concentration. The effect of the photoinitiator concentration on the conversion and polymerization rate of HDDA and EPETA on irradiation with monochromatic UV light is shown in Fig. 3 and 4. An increase of the initiator content strongly accelerates the photopolymerization of the monomer mixture, and the maximum of the polymerization rate is achieved earlier.

3 and 4. An increase of the initiator content strongly accelerates the photopolymerization of the monomer mixture, and the maximum of the polymerization rate is achieved earlier. In contrast, the induction period is hardly influenced except for the lowest concentration. The polymerization rate Rp in a light-induced reaction depends (i) on the number of absorbed photons, (ii) the efficiency to use the deposited energy for the generation of radicals, and (iii) the reactivity of those radicals with the monomer and the growing chain.

It can be calculated according to,R = pk p 0.5kt[]( M φ Ii abs)0.5Where, kp and kt and the rate constants of chain propagation and termination respectively, [M] is the monomer mixture concentration, φi is the initiation quantum yield of the photoinitiator and Iabs is the absorbed intensity in the sample (Scherzer and Decker 1999).The absorbed intensity is related to the intensity of the incident radiation (I0) by, ) ( I = I 1-exp -2.303A abs()0A is the absorbance of the sample which is given by Lambert-Beer's law as, A= ε[PI]d (7) where ε is the molar extinction coefficient at the wavelength of the light, [PI] is the photoinitiator concentration, and d is the thickness of the sample. Using Eqs. (6) and (7), the polymerization rate can be expressed as,R = pk p 0.5kt[((( ] M φ I 1-exp -2.303ε PI d][i 0)))Eq. (8) describes a relationship between the polymerization rate and the photoinitiator concentration.

(8) describes a relationship between the polymerization rate and the photoinitiator concentration. However, this relation naturally depends on [M] even when [PI] is the only(5)(6)(8)experimental parameter which is varied. Therefore, a plot of Rp vs photoinitiator concentration has to be made at equal acrylate conversion. In Fig. 4, the polymerization rate was plotted vs. conversion. It can be clearly seen that the maximum polymerization rate Rp,max occurs at the same conversion (i.e. at about 30%) at all photoinitiator levels studied(Scherzer and Decker 1999). Accordingly, Rp,max can be plotted vs. the square root of the fraction of the absorbed light ( Fig. 5). It is obvious that Rp,max is proportional to (1-exp(-2.303 [PI]d ))0.5 as predicted by Eq. (8). Remarkably, no significant deviation from linearity is observed at higher photoinitiator concentrations. Such behavior could be expected since the absorbance has to be small for validity of Eq. (8) (Scherzer 2002).

Dependence of Rpmax on the fraction of absorbed energy for varying photoinitiator concentrations,3530n o25i s r e v n o c2015%conversion vs PI concentration%105000.10.20.30.40.50.60.7PI IR 184 concentration( wt %)Fig6. Variation of conversion with Photoinitiator concentration for fastest processing time on LAMP.Effects of Filler content As apparent in Fig. 7 and 9 conversion and rate of polymerization progressively decreased with increasing filler-loading at a given energy dose. This is in agreement with (Alvarez-Gayosso,Barcelo-Santana et al. 2004; Atai and Watts 2006). At higher filler loading, the mobility of resin- monomers can be restricted on the surface of the fillers, leading to a decrease in molecular and radical mobility and resulting in lower conversion. However the relative conversion over the range of 10% to 80% filler loading was only about 20 to 25%.

However the relative conversion over the range of 10% to 80% filler loading was only about 20 to 25%. Fig 10 further confirms the insignificance of conversion over varying filler loading at 60ms of exposure time on LAMP machine. Thus, it seems that the influence of filler on the conversion and reaction kinetics of the composite is more related to filler particle-size and surface-area than filler loading (Atai, Watts et al. 2005). As in the case of varying photoinitiator concentration, gel point occurs within the same range of 40 to 60% conversion irrespective of the filler content (Fig 8).90n o i s r e v n o c %8070605040302010010% FL20% FL30%FL40%FL50%FL60%FL70%FL0.911.11.21.31.41.580%FLEnergy Dose(kJ/m2)Fig7. Effect of filler loading on conversion0.60.510% FL) s / l / l0.430% FLo m ( p R0.30.20.150%70%FL80% FL0020406080100% conversionFig8.

However from Fig 11 and 12, higher conversion results in increased volumetric shrinkage strain. Hence it is desired to have lower filler content at the expense of increased conversion. This expense can again be compensated by 2 different means. The first choice is to increase the photo initiator concentration and this again leads to the problem of higher shrinkage strain (Fig 11) and hence is ignored, while the second recommendation is to have an excess of monomer molecules in order to completely utilize the availability of initiating species. This procedure is expected to ensure lower shrinkage strain without compromising on the desired conversion. Further from Eq 4 it can be seen that the slopes of the curve in Fig 11 and 12 would remain constant for a given class of monomers. The shrinkage strain for 0.43% IR 184 at 60% filler loading which is presently used in LAMP machine is found to vary from 6% to 13% over a conversion range of 30 % to 68%.

The shrinkage strain for 0.43% IR 184 at 60% filler loading which is presently used in LAMP machine is found to vary from 6% to 13% over a conversion range of 30 % to 68%. Further thistechnique can be used to monitor conversion over several different layers as the build progresses and can be used to determine the cumulative shrinkage strain. This data when used along with the slicing algorithms can accommodate for the shrinkage strain resulting in closer conformation of the desired part.18)% ( n16i a r t14S e g a k n i r h s121080.2 % PI 1840.35% PI 1840.40% PI 1840.45% PI 184c i r t e m u640.5%PI 184lo V200102030405060708090Degree of Conversion (% )Fig 11.

E CHARACTERIZATION OF THE PERFORMANCE OF A NEW POWDERFEEDER FOR LASER BASED ADDITIVE MANUFACTURINGH. Mei, M. Valant, D. Hu, R. KovacevicResearch Center for Advanced Manufacturing Southern Methodist University 1500 International Parkway, Suite 100, Richardson, TX, 75081AbstractLaser-based additive manufacturing (LBAM) requires precise control over the metal, ceramic, or carbide powder added to the molten pool. The feeding rate of the powder must be very consistent, and it must respond rapidly to commands to change the feeding rate. LBAM also requires feeding rates as low as one gram per minute. Currently, commercially available powder feeders are optimized for such tasks as feeding powder to thermal spraying processes, which generally require a much higher feeding rate than LBAM, and can usually tolerate much more variation in the feeding rate. These powder feeders are therefore not suitable for the LBAM process.

These powder feeders are therefore not suitable for the LBAM process. The Research Center for Advanced Manufacturing at Southern Methodist University has designed and built a new powder feeder capable of consistent, repeatable powder delivery at extremely small flow rates. The powder feeder is regulated by a weight-based control system, which provides real-time measurement of the mass remaining in the feeder as powder is transferred to the powder nozzles. The powder feeder has been fully characterized to obtain correlations between the input parameters, powder type and the resulting mass flow rates. The powder nozzles at the laser head have also been characterized. The nozzle angle, standoff height, and carrier gas flow rate have each been optimized experimentally to maximize the concentration of powder arriving at the molten pool created by the laser beam, as detected using a sheet of He-Ne laser light and a coaxial vision system.

The powder delivery efficiency of the system has been thus maximized, increasing both the deposition rate and the quality of the deposited material.Keywords: laser based additive manufacturing, powder feeder, flow rate, control1. IntroductionLaser based additive manufacturing (LBAM) is a novel manufacturing technology which uses a laser beam to melt the additive material to produce functional parts with freeform near net-shape and the required mechanical properties. LBAM can produce relatively complex three-dimensional structures like inner cavities, which are normally difficult or448impossible for traditional manufacturing technologies based on material removal. LBAM is also able to change the chemical composition at various positions on the parts in accordance with the property requirements and to produce functionally gradient materials (FGMs).

LBAM is also able to change the chemical composition at various positions on the parts in accordance with the property requirements and to produce functionally gradient materials (FGMs). These unique features make LBAM a suitable technology for the manufacturing and repair of high value industrial components in the tooling industry and in the aerospace and aircraft industries [1-4].The most commonly used materials in LBAM are ceramic and metal powders. To produce a part with high dimensional accuracy and the desired composition distribution, LBAM requires precise control over the metal, ceramic or carbide powder added to the molten pool created by the laser beam. The feeding rate of the powder must be very consistent, and it must respond rapidly to commands to change the feeding rate.

The feeding rate of the powder must be very consistent, and it must respond rapidly to commands to change the feeding rate. LBAM also requires feeding rates as low as one gram per minute.Currently, the commercially available powder feeders are optimized for such tasks as feeding powder to thermal spraying processes, which generally require a much higher feeding rate than LBAM, and can usually tolerate much more variation in the feeding rate. These powder feeders are therefore not suitable for the LBAM process. In recent years, several attempts have been made to develop new powder feeders for laser deposition technologies [5-9]. B. Gruenenwald, etc. [5] designed a powder feeding system for the requirements of laser surface treatment. This powder feeder used on-line measurement and control of the powder mass flow rate and was characterized by feed rates of between 1 and 150g/min and an accuracy of better than 5%. X. Yang, etc.

This powder feeder used on-line measurement and control of the powder mass flow rate and was characterized by feed rates of between 1 and 150g/min and an accuracy of better than 5%. X. Yang, etc. [6-7] developed a powder feeder for larger-area laser cladding. This powder feeder used no carrier gas, and achieved a powder flow rate of 0.5-200g/min and a variation of flow rate of less than 2%. However, the two systems described above were mainly designed for the application of high power laser cladding and could feed only a single kind of powder at one time. They could not be used in the applications where multiple powders are needed, e.g. in the production of functionally graded materials. The literature [8-9] described a powder feeder that could deliver three kinds of powder at the same time. The system utilized a load cell based on a semi-conductor strain gauge and a photodetector to measure the powder flow rate, and used close-loop control over the triple-hopper powder feeder.

The system utilized a load cell based on a semi-conductor strain gauge and a photodetector to measure the powder flow rate, and used close-loop control over the triple-hopper powder feeder. This powder feeding system exhibited the flexibility, efficiency, and the improvement of accuracy of multiple powder delivery. However, the measuring system only measured the total flow rate of the mixed powder and could not precisely control the composition of each individual powder. The composition of each individual powder is actually very important for the production of functionally graded materials. Therefore, these powder feeders are not the best fit for the LBAM process.The aim of the present work is to develop a powder feeder suitable for the LBAM process. In the following part, the principle and the performance of this newly developed powder feeder are described.

In the following part, the principle and the performance of this newly developed powder feeder are described. The third part gives the results of laser deposition experiments using the newly developed powder feeders, followed by the conclusions in part 4.4492. Development of Powder Feeder2.1 Principle of powder feederThe powder feeder consists of a powder storage hopper, a dosing disk coupled to a DC stepper motor, and a pneumatic powder delivery subsystem. Under the force of gravity, the powder from the hopper flows through a funnel and then falls onto the rotating horizontal dosing disk. The carrier gas, normally Argon, carries the powder into the outlet tube and transports it to the laser head. Fig.

The carrier gas, normally Argon, carries the powder into the outlet tube and transports it to the laser head. Fig. 1 shows two such new powder feeders used for LBAM, especially for the production of Functionally Graded Materials (FGMs).The powder flow rate depends on the diameter of the funnel hole, the gap between the funnel tip and the dosing disc, and the rotating speed of the dosing disc. Changing the funnel diameter and the funnel-disc gap gives different operating ranges of the powder feed rate. For a fixed funnel diameter and funnel-disc gap, the powder flow rate is nearly proportional to the speed of the DC stepper motor.Fig 1: New powder feeders used for production of FGMs(a)(b)Figure 2. Powder feeder control system.450To fabricate functionally graded materials, multiple powder feeders are used. Control systems are studied to achieve the control of multiple powder feeders.

Powder feeder control system.450To fabricate functionally graded materials, multiple powder feeders are used. Control systems are studied to achieve the control of multiple powder feeders. A prototype controller is created using LabView and the corresponding digital I/O module is installed on a personal computer (PC) (Figure 2a). The controller generates pulse sequences with the desired step frequencies to control the step motors on the powder feeders. By receiving commands from the motion system through the PLC output, the desired composition of the FGM at the specific position can be obtained by setting the step frequency of each step motor. Based on the success of the prototype controller, an embedded powder feeder controller is also designed (Figure 2b). Each controller is equipped with an embedded microcontroller, LED display, keypad, and Ethernet and RS485 connections, and controls one powder feeder.

Each controller is equipped with an embedded microcontroller, LED display, keypad, and Ethernet and RS485 connections, and controls one powder feeder. The controller and the powder feeder are integrated into a standalone powder feeding system that can work independently if desired. The controller interfaces with the operator through the LED display and 16-key keypad. The desired delivery rate can be set from the keypad, and the working status of the controller is shown on the LED display. Several such powder feeding systems can be linked to the central control PC through Ethernet or RS485 serial connections to form a multiple-powder feeding system. A system control software support TCP/IP and serial port communication protocols runs on the central PC. The desired delivery rate of each kind of powder is controlled by the central PC in real time by sending commands through Ethernet or RS485 serial connections. Up to four powder feeders can be controlled simultaneously by the central PC.

Up to four powder feeders can be controlled simultaneously by the central PC. The embedded powder feeder controller together with the powder feeding machine delivers a low cost and modularized powder feeding system ready for commercialization.2.2 Calibration of Powder FeederThere are a lot of parameters which affect the powder feed rate, such as the powder type, the particle size, the diameter of the funnel hole, the funnel-disc gap and the motor speed. For a specific operating condition, a calibration procedure is needed to determine the numerical relation between the motor speed and the powder feed rate. An automatic calibration program is made using the C++ language. During the powder feeding process, the weight of the powder feeder is continuously recorded using an accurate electronic balance. The actual powder mass flow rate is calculated as weight loss per unit time. Fig 3 shows the powder feed rate as a function of motor speed.

The actual powder mass flow rate is calculated as weight loss per unit time. Fig 3 shows the powder feed rate as a function of motor speed. The H13 powder used in the calibration has a mesh size of –100/+325. The mesh size of the Tungsten Carbide powder is –80/+200. As can be seen in the figure, a very good linear relation exists between the powder flow rate and the motor speed. For H13 powder, the powder flow rate could be as low as 0.5g/min.451222018H13n m / g ( e t a ri16141210W C2w o8F64200 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25Motor Speed (rpm)Fig 3: Powder flow rate as function of motor speed2.3 Powder Delivery FeaturesThe powder delivery features, such as the concentration mode, the powder distribution and the particle velocity are essential for laser deposition quality and powder efficiency [10- 11].

If the powder particle density in the laser path is too high, a large portion of laser power will be absorbed by the flying powder particles, and less power will remain for the creation of the molten pool. As a result, the deposited cladding will be poorly bonded to the substrate [10]. On the other hand, if the powder particle density and the particle speed are very low, the powder particles will be burned due to excessive heating. Also, the powder catchment efficiency can be improved if the powder stream is made to be more concentrated [11].Powder delivery features depend on the performance of the powder feeder as well as the nozzle structure of the laser head. For the LBAM unit at SMU, a four-nozzle coaxial laser head is used. The four nozzles symmetrically surround the laser beam. Because the powder flow rate of this newly developed powder feeder can be as low as 1g/min, the diameter of nozzles is as small as 0.6mm.

The four nozzles symmetrically surround the laser beam. Because the powder flow rate of this newly developed powder feeder can be as low as 1g/min, the diameter of nozzles is as small as 0.6mm. A smaller nozzle diameter results in a more concentrated powder stream and a higher powder efficiency. Moreover, the powder flow rate is independent of the carrier gas pressure. So the carrier gas pressure used by the powder feeder can be adjusted to optimize the particle speed.An optical imaging method is used to measure the powder stream structure and the powder distribution within that stream. The experimental setup is shown in Fig. 4. A horizontally placed He-Ne laser sheet is used to illuminate the powder stream. When the powder particles pass through the laser sheet, the red He-Ne laser light is scattered. The higher the powder particle density becomes, the brighter the luminosity on the cross section of the powder stream.

The higher the powder particle density becomes, the brighter the luminosity on the cross section of the powder stream. So the luminance distribution on the cross section of the powder stream indicates the powder distribution pattern on that cross section. On the top of the laser head, a CCD camera is coaxially installed to take the image of the illuminated cross section of the powder stream. The images taken at various standoff heights from the nozzle tip reveal the spatial structure and the concentration mode of the powder stream.452Fig. 5 shows the luminance intensity distribution at various standoff heights for nozzles with tilting angles of 45° and 60°. It can be seen that the most concentrated section of the powder stream (focus of powder stream) lies at the standoff height of 0.10 inches under the nozzle tip for 45° nozzles or at 0.15 inches for 60° nozzles.

When the standoff distance between the nozzle tip and the substrate is set to these values, the highest powder efficiency is achieved. It also can be seen from Fig. 6 that the area for the focus of powder stream for 60° nozzles is smaller than for 45° nozzles, while the brightness is greater than for 45° nozzles. This indicates that the concentration degree for the 60° nozzles is larger than for 45° nozzles. Therefore the resulting powder efficiency of 60° nozzles will be higher, as demonstrated by various laser deposition experiments at RCAM. In addition, the working standoff for 60° nozzles is larger than that of 45° nozzles, which greatly reduces the possibility of interference with the nozzles by the scatter from the laser-substrate interaction zone.

Therefore, in our LBAM system, the 60° tilting angle for the nozzles is preferred.CCD CameraIrisNd: YAG LaserNd: YAG FilterPartially Reflective MirrorLBAM Laser HeadPowder Delivery NozzlePowderHe-Ne laserFig 4: Setup to record the powder streamS0.04 in0.06 in0.08 in0.10 in0.12 in0.14 in45°PS0.06 in0.09 in0.12 in0.15 in0.18 in0.21 in60°P(S = Standoff; P = Powder stream pattern)Fig 5: Powder stream pattern at various nozzle standoffs (nozzle angle 45° and 60°)3. Laser Deposition Experiments4530.16 in0.24 inThe schematic diagram of the LBAM unit used in this research is shown in Fig.6. Two powder feeders described above are built and integrated in the LBAM unit. Two kinds of materials (H13 and WC powders) are utilized. Single-layer-wide walls with functionally gradient material are made.The wall is produced by a stepwise deposition process with layers deposited on the tops of previous layers.

Single-layer-wide walls with functionally gradient material are made.The wall is produced by a stepwise deposition process with layers deposited on the tops of previous layers. Starting with the pure H13 tool steel powder, the amount of WC-NiBSi ceramet in the powder mixture is increased steadily to generate materials with a compositional gradient. Fig. 7 shows the designed distribution of different constituents in the as-deposited FGM wall.

6: Schematic diagram of LBAM unit100nis t n e u t i t s n o ct n e r e f f i d%. tw M G F d e n g,908070605040f os t n e t n o Ci s e d302010H13 tool steel WC particle WC(NiSiB) ceramet002468101214Distance from the top surface, m mFig. 7: Designed distribution of different constituents in the as-deposited FGM454Fig. 8 shows the deposited single-layer-wide wall with functionally graded layers. In the image depicting the surface morphology (Fig. 8(a)), the color change caused by composition change is obvious. The polished and etched cross-section (Fig. 8(b)) of the sample also clearly shows the change of the composition. At the bottom of the wall, the wall appears dark, showing that this part is made of pure H13 tool steel, because the pure H13 material is easily to be etched.

At the bottom of the wall, the wall appears dark, showing that this part is made of pure H13 tool steel, because the pure H13 material is easily to be etched. From the bottom to the top of the wall, the color of the cross-section becomes brighter, indicating that the composition of WC is increasing, since the WC particles are very hard and chemically resistant, and thus appear shiny. Fig. 9 shows the enlarged morphology of the WC particles at the cross-section of the graded WC/H13 wall. It is evident in this figure that the amount of WC particles increases from the bottom to the top of the wall, corresponding to the designed composition gradient.Fig. 8 As-deposited single layer wide wall with functionally graded layers (a) Surface morphology; (b) Cross-sectionFig.

Each powder feeder is equipped with a weight-based controller and is capable of consistent, repeatable powder delivery at relatively small flow rates. Up to 4 such newly developed powder feeders could be integrated into the LBAM unit and precise control over each individual powder delivery can be achieved to meet the requirements of the production of Functionally Gradient Materials (FGMs) by LBAM. The nozzle angle, standoff height, and carrier gas pressure have each been optimized experimentally to maximize the concentration of powder arriving at the molten pool created by the laser beam, as detected using a sheet of He-Ne laser light and a coaxial vision system. The powder delivery efficiency of the system was thus maximized, increasing both the deposition rate and the quality of the deposited material.5. AcknowledgementThis work was financially supported by THECB, Grants 003613-0022-1999 and 03613- 0016-2001, NSF, Grants No. DM1-9732848 and DM1-9809198, and the U.S.

ARACTERIZATION OF THIN WALLED Ti-6Al-4VCOMPONENTS PRODUCED VIA ELECTRON BEAM MELTINGDenis Cormier, Harvey West, Ola Harrysson, and Kyle Knowlson North Carolina State University Department of Industrial Engineering Raleigh, NC 27695-7906Reviewed, accepted August 4, 2004Abstract Direct-metal energy beam SFF processes typically produce layers by scanning the contours and then filling in the area within the contour. Process parameters used to solidify contours are often different from those for fill areas. It is to be expected, therefore, that the contour and fill area regions will have different microstructures. This can have important ramifications for thin walled components such as biomedical implants whose slices have very little fill area. This paper characterizes the metallurgical differences in contour and fill areas in titanium components produced via Electron Beam Melting.

This paper characterizes the metallurgical differences in contour and fill areas in titanium components produced via Electron Beam Melting. The implications of these properties for thin walled components are described.Introduction Arcam's Electron Beam Melting (EBM) process (www.arcam.com) is a direct-metal freeform fabrication process that uses a 4.8 kW electron beam to selectively scan and melt layers of metal powder one on top of the next. Details of the process can be found in reference [1]. Titanium (Ti-6Al-4V) is an alloy that is receiving considerable interest for use with this process from industry. The EBM process was originally developed with the tool and die making industry in mind. Whereas the H13 tool steel components originally targeted with this process are typically large bulky parts, the titanium components built with the process tend to be thin walled in nature. For purposes of this paper, a thin walled part is a part with a thickness of approximately 14 mm.

For purposes of this paper, a thin walled part is a part with a thickness of approximately 14 mm. Representative thin-walled Ti-6Al-4V components built with this process include the turbocharger compressor wheel shown in Figure 1(a), the knee implant shown in Figure 1(b), and the bone plate shown in Figure 1(c). The vanes on the turbocharger wheel are approximately 1.2 mm thick, and the knee implant and bone plate each has an average thickness between 3-4 mm.Most beam-based freeform fabrication processes melt or otherwise solidify a raw material by scanning the contour of a given slice and then scanning the fill area inside of the contours. In the case of direct-metal freeform fabrication processes, it is not uncommon for different process settings to be used when solidifying contours versus fill area. This is done primarily for reasons of improving surface finish and appearance.

This is done primarily for reasons of improving surface finish and appearance. As the solidification rate of the metal powder can be influenced by differences in process settings (e.g. beam power and scan speed), it is possible for the contours and fill area to have substantially different microstructure and mechanical properties.440(a)(b)(c)Figure 1 – (a) Ti-6Al-4V Turbocharger Wheel; (b) Knee Implant; (c) Bone PlateFigure 2 shows the slice data for a relatively bulky part on the left and a thin walled part on the right. A typical melting cycle for the layer illustrated in Figure 2(a) would start by melting the inner and outer contours. For purposes of improving surface finish, it is not uncommon to melt two contours, with the second contour being offset inward from the first contour by a user- specified amount. After the contour is melted, the area shown by dashed lines is melted.

After the contour is melted, the area shown by dashed lines is melted. Although the system control software varies the beam current and scan speed to achieve the necessary heat input, the contours are often melted at approximately 100-200 mm/sec with a beam power of 300-600W. In contrast to the relatively low melting speed for contours, the fill region is typically melted with a scan speed in the 400-700 mm/sec range and a beam power in the 1,200-1,800W range. Although different melting speeds are used, the beam power is adjusted441so that the overall power density during melting is approximately the same for both contours and fill area.Figure 2(b) shows the slice data for a thin-walled part. Note that the entire melting cycle consists solely of melting a single contour since it is not possible to offset a second contour. Likewise, there is no fill area to be melted.

Note that the entire melting cycle consists solely of melting a single contour since it is not possible to offset a second contour. Likewise, there is no fill area to be melted. Since contours are melted more slowly in order to improve the surface finish and appearance of the part, lower melting speeds could potentially result in lower solidification rates. On the other hand, the speed with which a part looses heat following processing is largely determined by conduction between the part and the surrounding powder bed. The higher surface area-to volume ratio of thin-walled parts versus bulky parts would logically imply that thin-walled parts cool down more quickly than bulky parts.As Arcam's EBM process was originally targeted towards the fabrication of relatively bulky parts, there are no "build styles" specifically developed for the fabrication of thin-walled titanium parts whose slices consist primarily of contours.

The aim of this paper, therefore, is to examine and characterize the metallurgical differences between thin-walled and bulky components built via the EBM process using Arcam's default settings.(a)(b)Figure 2 - Bulk Versus Thin-Walled Melt GeometryExperimental Procedure Two Ti-6Al-4V components were fabricated using an Arcam EBM machine – one "bulky" part and one "thin-walled" part as shown in Figure 3. For microstructural examination, each part was sectioned perpendicular to the build direction using a water-cooled silicon carbide cut-off saw. Each specimen was then mounted in a room-temperature curing resin. After grinding with successively finer grits of emery paper, the samples were polished using 6 µm diamond paste, 1 µm alumina, and 0.3 µm alumina.

After grinding with successively finer grits of emery paper, the samples were polished using 6 µm diamond paste, 1 µm alumina, and 0.3 µm alumina. Prior to metallographic examination, the polished surfaces were etched with Kroll’s reagent.442Figure 3 - Bulky (L) and Thin Walled (R) PartsResults and Discussion When heated, titanium undergoes an allotropic phase transformation at 882°C, changing from close packed hexagonal α-Ti to body centered cubic β-Ti. Therefore, the as-cast microstructure of Ti-6Al-4V primarily consists of acicular α-Ti that has transformed from the grains of β-Ti that solidified initially. In addition, α-Ti is also present at the prior grain boundaries of the β-Ti [2]. This same microstructure is observed in the parts produced via the EBM process, as shown in Figures 4 and 5.

In addition, α-Ti is also present at the prior grain boundaries of the β-Ti [2]. This same microstructure is observed in the parts produced via the EBM process, as shown in Figures 4 and 5. Of particular significance from these micrographs is the fact that the α-Ti needles produced in the thin wall specimen are approximately one third the size of those produced in the bulk part. Microhardness measurements taken on both samples were within the range of HV 285-300 (HRC 28-30).200 µmFigure 4 - Thin-Walled Microstructure at 100X Magnification443200 µmFigure 5 - Bulk Part Microstructure at 100X MagnificationExamination of the prior β-Ti grain boundaries revealed that epitaxial growth occurred between layers. In Figures 4-6, the build direction is from the bottom to the top of the micrograph. Given that each build layer is nominally 150 µm thick, each micrograph spans multiple horizontal layers, with Figure 6 displaying 10-11 horizontal layers.

Given that each build layer is nominally 150 µm thick, each micrograph spans multiple horizontal layers, with Figure 6 displaying 10-11 horizontal layers. Note that there is no observable delineation between successive layers in either the bulk or thin-walled parts and that the prior β-Ti grain boundaries are continuous through the layers. The growth of β-Ti grains is particularly evident in Figure 6. This implies that the part temperature is maintained above the β transus, and that solidification of the melted powder occurs by growth of β-Ti directly onto the “seed” grains of the previous layer. Control and characterization of this mode of solidification will be the subject of further investigations.During processing, the EBM process control algorithm is designed to keep the part being fabricated at an elevated temperature in order to reduce thermal stresses.

While thermal imaging techniques may be used to measure surface temperatures at the top of a part during the build [3], it is quite difficult to measure temperature within the body of a part during a build due to the nature of the process. However, each part with this process is built up from a steel base plate, and a thermocouple is attached to the underside of the base plate. A base plate temperature of 850°C was maintained during the builds that produced all parts shown in this paper. There is obviously an increasing temperature gradient from the underside of the start plate to the very top surface of444the part that extends above the melting temperature of the alloy. With a start plate temperature of 850°C, the bulk of the part is above the β transus for the duration of the build. The α-Ti needles therefore do not form until the entire part is cooled below the β transus following the completion of electron beam melting. The process parameters (e.g.

The α-Ti needles therefore do not form until the entire part is cooled below the β transus following the completion of electron beam melting. The process parameters (e.g. beam current and scan speed) had relatively little influence on the size of the α-Ti needles. Instead, the cooling rate and resulting microstructure was largely dominated by the part geometry in these cases. In a vacuum environment, heat loss through conduction with the surrounding powder bed is the primary cooling mechanism, and thin walled parts will cool down more rapidly than bulky parts for a given input power density. The thin walled part represented in Figure 4 was melted at a much slower speed than the bulky part shown in Figure 5, yet it has a much more refined microstructure.

The thin walled part represented in Figure 4 was melted at a much slower speed than the bulky part shown in Figure 5, yet it has a much more refined microstructure. Within the range of values recommended by the process control algorithm, it was clear that part geometry had more to do with microstructure than processing parameters.400 µmFigure 6 - Micrograph of Bulk Part at 50X Magnification445Conclusions and Future Directions Each layer that is melted in the EBM process involves both contours and fill area. Contours are melted at roughly 25% of the speed of fill area in order to improve surface finish. Because each layer of a thin wall part consists primarily of contours rather than fill area, thin walled parts are built primarily using the slower scan speeds associated with contours. Conversely, bulky parts whose layers are primarily fill area are built using much higher scanning speeds.

Conversely, bulky parts whose layers are primarily fill area are built using much higher scanning speeds. For the thin wall components such as those shown in Figure 1, all surfaces will be subject to light finish machining. Any improvement in surface finish resulting from slower scan speeds of the contours is largely for cosmetic purposes and will not affect the final form of the part following finish machining. It is therefore worth investigating whether or not a special build style can be used to optimize the microstructure and properties of thin wall parts.Using the generic recommended processing parameters, experiments determined that the refined microstructure of the thin wall part was due primarily to the effect of geometry on cooling rate. While there was no significant difference in microhardness, the finer microstructure of the thin wall part is preferred in fatigue applications where fracture toughness is important [4] [5].

While there was no significant difference in microhardness, the finer microstructure of the thin wall part is preferred in fatigue applications where fracture toughness is important [4] [5]. The turbocharger and bone implant applications shown in Figure 1 certainly fall into this category.The growth of columnar grains demonstrated in these experiments could give rise to EBM parts with tailored anisotropic microstructures. Titanium Aluminide (TiAl) is an alloy of particular interest as a lighter weight substitution for nickel-based superalloys in elevated temperature applications. If the same columnar growth behavior can be induced with TiAl, then the result would be similar to directional solidification. In that case, components could be oriented within the build chamber to maximize benefits of the columnar grain structure.The presence of columnar grain growth spanning multiple layers suggests that the temperature of the parts were held above the β transus throughout the build.

During processing, the control algorithm uses the electron beam to deliver extra heat to both the part and the surrounding powder bed in order to reduce thermal stresses. It is not strictly necessary to maintain the elevated temperature above the β transus. Future experiments will therefore investigate the effects of maintaining a body temperature that is below the β transus to see if a more rapid transition from β-Ti to α-Ti will take place during processing. Since the size of the β- Ti grains impose an upper limit on the size of the α-Ti needles, it is conceivable that increasing the scanning speed and beam current to those levels currently used for fill melting will produce a even finer α-Ti grain structure. For components such as those shown in Figure 1 that undergo cyclic stresses, this further refinement of microstructure could prove to be quite beneficial in terms of fatigue performance.

For components such as those shown in Figure 1 that undergo cyclic stresses, this further refinement of microstructure could prove to be quite beneficial in terms of fatigue performance. It is recognized that by lowering the target body temperature, the potential for tradeoffs between microstructural refinement and residual stresses will be considered as part of the studies.increased residual stresses and warping exists. The446References[1] Cormier, D., Harrysson, O., and West, H. (2004) "Characterization of H13 steel produced via electron beam melting" Rapid Prototyping Journal, V10N1, pp. 35-41.[2] Metals Handbook 9th Edition, Volume 9: Metallography and Microstructures, American Society for Metals, 1985.[3] Hofmeister, W., Wert, M., Smugeresky, J., Philliber, J., Griffith, M., and Ensz, M. (1999) "Investigating Solidification with the Laser-Engineering Net Shaping (LENS) Process," JOM, Vol. 51, No.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCharacterizing Interfacial Bonds in Hybrid Metal AM StructuresJohn Linn, Jason M. Weaver, Michael P. Miles, Yuri Hovanski Brigham Young UniversityRobert Smith Qualified Rapid ProductsAbstractThe capabilities of various metal Additive Manufacturing (AM) processes, such as Powder Bed Fusion - Laser (PBF-L) are increasing such that it is becoming ever more common to use them in industrial applications. The ability to print atop a substrate broadens that scope of applications. There is ongoing research regarding the mechanical properties of additively processed materials, but not much regarding the interaction between additive material and its substrate. An understanding of the mechanical and performance properties of the AM/substrate interface is imperative.

An understanding of the mechanical and performance properties of the AM/substrate interface is imperative. This paper describes a study of the strength properties of AM/substrate interfaces, with respect to torsion and tension, and compares them to their fully wrought and fully additive counterparts.IntroductionThere is an overwhelming amount of waste in the metal tooling industry. It isgenerally not worth the time or money to repair a metal tool, as it is cheaper to simply replace it. This is especially true when the tool needing replaced is either very simple, thus being easy to mass produce, or very intricate, making it difficult to repair. There are various methods used to repair metal products. Metal filled epoxies, mechanical fasteners, and welding techniques are common solutions to reworking parts for continued use. Reworked tools and products tend to have decreased mechanical properties that may influence their performance.

Reworked tools and products tend to have decreased mechanical properties that may influence their performance. In many industries the safety standards require mechanical properties that are unattainable through these common reworking processes. A remanufacturing process that provides similar or better mechanical properties than the original part, while also being time and cost effective, is a current need in the engineering and manufacturing fields.Additive manufacturing (AM) is one viable solution to this need. The last decade has seen an explosion of growth in metal additive manufacturing processes [1]. The quality and consistency of these processes [2] are reaching sufficient levels to support not only rough prototyping, but repair, tool design, and even full-scale manufacturing. There are numerous benefits to producing metal parts additively, rather than by other conventional casting and machining methods.

There are numerous benefits to producing metal parts additively, rather than by other conventional casting and machining methods. Time, money, and resources can be preserved, while also dramatically increasing design flexibility [3]. Custom parts can be fabricated rapidly, from design phase to solid part, in a matter of hours. This research focuses on the remanufacturing capabilities of AM.Utilizing AM as a remanufacturing process, many tools and products could berepaired rather than replaced, saving both time and money [4]. New material and features can be printed directly onto existing parts and structures, allowing the flexibility of AM without the1191relatively high cost of printing a new part from scratch.

New material and features can be printed directly onto existing parts and structures, allowing the flexibility of AM without the1191relatively high cost of printing a new part from scratch. However, the mechanical properties of the additively manufactured features and their interface to the substrate material must be sufficient to meet or exceed the requirements of the original part in order for this process to be viable.This paper describes the material properties testing of samples produced by printing stainless steel material onto a substrate of the same material, with a focus on the properties at and near the interface. There has been very little research focusing on the interfacial bond between a metal AM part and a substrate, although it has been touched on [5].

There has been very little research focusing on the interfacial bond between a metal AM part and a substrate, although it has been touched on [5]. Validating metal AM as a reliable process for repair of metal tools will have massive implications on many industries worldwide.MethodologyIn order to gain a better understanding of the mechanical properties and capabilities of the interface between AM material and a substrates, samples of SS316L stainless steel were prepared for torsion and tensile testing. Torsion and tensile tests were performed on three different categories: fully wrought specimens, fully additive specimens, and hybrid (AM/substrate) specimens. Both torsion and tensile specimen designs are according to ASTM E8 specifications.

Both torsion and tensile specimen designs are according to ASTM E8 specifications. The methods for testing these pieces allowed for both torsion and tensile tests to be run on the same Instron machine by simply altering the test method and switching the jaws.Sample Fabrication:The specimens were prepared from three blocks of material (Figure 1):One block wrought SS316L (0.5" x 2.25" x 3.0")One block AM SS316L (0.5" x 2.25" x 3.0")One hybrid block consisting of a 0.5" x 2.25" x 1.5" block wrought SS316L withadditional SS316L of the same dimension printed on top (resulting in a block the same dimension as the other two blocks)The AM portions of the test specimens were deposited using a ConceptLaser PBF-L typemachine, using SS316L stainless steel and a laser power of 49 J/mm3.Figure 1.

Wrought, AM, and hybrid blocks of SS316L stainless steel1192From these blocks, the dimensions in Figures 2 and 3 were used to prepare the followingspecimens (Figure 4):3 torsion test specimens of wrought SS316L3 torsion test specimens of AM SS316L3 torsion test specimens of hybrid AM/wrought SS316L3 tension test specimens of wrought SS316L3 tension test specimens of AM SS316L3 tension test specimens of hybrid AM/wrought SS316LFigure 2. Torsion test specimen (all dimensions in inches)Figure 3. Tension test specimen (all dimensions in inches)1193Figure 4. Torsion and tensile test specimens of hybrid AM/wrought SS316LTorsion Testing:Each torsion test specimen was placed in an Instron Tensile Tester machine (Figure 5)with a 10,000 lb-in torque load capacity. The first sample was loaded in torsion through 45 degrees of rotation. All subsequent samples were submitted to 90 degrees of motion (registered in the data as from -45 to +45 degrees).

The first sample was loaded in torsion through 45 degrees of rotation. All subsequent samples were submitted to 90 degrees of motion (registered in the data as from -45 to +45 degrees). Wave Matrix software was used to perform the experiments consistently and gather pertinent data (torque load vs. degrees twisted) to aid in making observations and drawing conclusions. Pictures were taken of the test specimens before and after testing in order to more accurately conclude what occurred.Tensile Testing:Each tensile test specimen was placed in an Instron Tensile Tester machine with a 20,000lb. tensile load capacity. The samples were loaded in quasi-static tension until ultimate failure. Wave Matrix software was used to perform the experiments consistently and gather pertinent data (tensile load vs. elongation torque load vs. time) to aid in making observations and drawing conclusions.

elongation torque load vs. time) to aid in making observations and drawing conclusions. Pictures were taken of the test specimens before and after testing in order to more accurately conclude what occurred.Figure 5. Instron machine used for torsion/tensile testing1194ResultsTorsion Testing:The first specimen tested, Wrought SS 1, was subjected to 45 degrees of rotation, thelimit of the Instron machine when starting from zero degrees. As shown in Figure 6, the material yielded, but continued to plastically deform without full failure. Because of this, the second specimen, Wrought SS 2, was subjected to 90 degrees of rotation, from -45 to +45 degrees on the machine, with similar results. The third wrought specimen was tested repeatedly by twisting the sample through 90 degrees, refixturing it in the machine, and twisting it again. After multiple full rotations, the specimen continued to plastically deform without failure.

After multiple full rotations, the specimen continued to plastically deform without failure. Data from Wrought SS 3 is not included below, due to the deviation from the test procedure. All subsequent specimens were subjected to 90 degrees of rotation like Wrought SS 2. Results of these tests are shown in Figures 6-8.Tensile Testing:All nine tensile test specimens were tested until failure according to the test procedure inthe previous section. Results of these tests are shown in Figures 9-11.1195Wrought SS 1Wrought SS 2e u q r o T n i - bl250 200 150 100 50 0e u q r o T n i - bl250 200 150 100 50 050050500Degrees RotationDegrees RotationFigure 6. Torque vs. rotation for wrought torsion specimensAdditive SS 1Additive SS 2e u q r o T n i - bl250 200 150 100 50 0e u q r o T n i - bl250 200 150 100 50 05005050050Degrees RotationDegrees RotationFigure 7. Torque vs.

Hybrid torsion specimen showing yieldingTensile Testing:Table 1 also shows the average tensile yield and ultimate strength for each set of specimens. As with the torsion tests, the AM specimens exhibited significantly higher strength than the wrought specimens. The hybrid specimens exhibited average yield strength of 40,000 psi (276 MPa) and ultimate strength of 75,000 psi (517 MPa), which is similar to the strength shown by the wrought samples. Again as with the torsion tests, visual inspection confirmed that the hybrid specimens all failed within the wrought material, well away from the interface between AM and wrought (Figure 13). All other specimens, both wrought and fully AM, yielded near the midpoint of the neck.WroughtAdditiveYieldingInterfaceFigure 13.

All other specimens, both wrought and fully AM, yielded near the midpoint of the neck.WroughtAdditiveYieldingInterfaceFigure 13. Hybrid tensile specimen showing yielding and fracturePredicted Strengths at the AM/Wrought Interface:Because all hybrid specimens yielded within the wrought material, not at the interfacewith the AM material, material properties of the interfacial bond could not be directly measured. However, as there appeared to be little to no yielding in this region of the specimens, we can put a lower bound of the interfacial strength at the failure of the wrought material. Thus in Table 1,1199the torsional and tensile yield strengths listed can be considered minim strengths, with the actual strengths at the interfaces considerably higher.

Thus in Table 1,1199the torsional and tensile yield strengths listed can be considered minim strengths, with the actual strengths at the interfaces considerably higher. It is unknown from these experiments whether the interfacial strength would be similar to the strength of the AM material or whether it would be somewhere between the two extremes.ConclusionFrom these eighteen tests, there are a few valuable conclusions that can be made:First, the AM specimens of SS316L stainless steel all had much higher yield and ultimatestrengths than the wrought specimens. There are significant processing differences between wrought metal and the micro-welding of laser-powered powder bed fusion [6]. It is expected that such differences will lead to differences in the two materials’ microstructure as well. We hypothesize that the micro-welding process of PBF results in smaller and more uniform grain boundaries than the working process.

We hypothesize that the micro-welding process of PBF results in smaller and more uniform grain boundaries than the working process. Future work for this study includes microscopic analysis of the grain structure of the two materials and the interface region between them.A second conclusion is that the interfacial bond of SS316L AM/substrate specimensoutperforms the original wrought material in both torsion and tension when using PBF-L as the additive process. The actual bond strength limits in torsion and tension were never reached. This is an encouraging result for applications such as repair and rework of rotary tools, molds, and other work pieces, as the newly deposited material and its bond to the wrought material appears to be predictably stronger in tension and torsion than the original part. Further characterization of the AM material and the interfacial bond in terms of impact hardness, ductility, bending, T-peel, etc.

Further characterization of the AM material and the interfacial bond in terms of impact hardness, ductility, bending, T-peel, etc. may be beneficial in fully understanding the tradeoffs in this area.Additional future work planned for this research area include performing this study using a different AM process (direct energy deposition), performing the study using PBF-L and one or more different alloys (such as M300 steel), and the previously mentioned analysis of grain microstructure before and after each of these tests.Works Cited1.2.3.4.5.Frazier, W.E., Metal Additive Manufacturing: A Review. Journal of Materials Engineering and Performance, 2014. 23(6): p. 1917-1928. Marco, G. and C. Bianca Maria, Process defects and in situ monitoring methods in metal powder bed fusion: a review. Measurement Science and Technology, 2017. 28(4): p. 044005. Atzeni, E. and A. Salmi, Economics of additive manufacturing for end-usable metal parts.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCHARACTERIZING INTERNAL POROSITY OF 3D-PRINTED FIBER REINFORCED MATERIALSFrye L. Mattingly1,2, Alan Franc2,3, Vlastamil Kunc1,2, Chad Duty1,2 1 University of Tennessee, Knoxville 2 Manufacturing Demonstration Facility, Oak Ridge National Laboratory 3Techmer PM, Clinton Tn.ABSTRACTAs the functional requirements for 3D printed parts become more demanding, the use of fiber reinforced materials in material extrusion printers is increasingly common. Although fiber- reinforced thermoplastics offer higher stiffness and strength, the internal volume of the extruded material often has a high degree of porosity which can negatively impact mechanical properties.

This research surveys the internal porosity present across a range of material extrusion additive manufacturing platforms, primarily those involving a single screw extruder, such as the Big Area Additive Manufacturing (BAAM) system. The porosity within the volume of an extruded bead was quantified through image analysis of cross sectional micrographs. The impact of extrusion rate, transient vs steady state flow, multiple hardware configurations, and material conditions were evaluated. Across the five systems studied porosities ranged from 0.1% to 18.4% with the greatest reductions in porosity coming from two systems that added a vent to the extruder barrel which lowered porosity 64% in one case and 98% in the other.1. INTRODUCTIONOne of the most ubiquitous problems within the field of polymer-based Additive Manufacturing (AM) is that of porosity or internal voids. When looking at a cross-section of a printed part, the beads of material look like bricks in a wall.

When looking at a cross-section of a printed part, the beads of material look like bricks in a wall. Most other work on porosity has been focused on the gaps or “mortar” between the beads or “bricks.” This study is instead focusing on characterizing the porosity within the beads themselves, effectively studying the quality of the “bricks.” These two types of porosity can be seen in Figure 1. The initial motivation for this study was that internal porosity of varying magnitudes had been observed for years within samples studied at Oak Ridge National Laboratory’s (ORNL) Manufacturing Demonstration Facility (MDF), but the source and variability of the porosity was not clearly understood. Several questions were posed that this study will address.

Several questions were posed that this study will address. How widespread is porosity in 3D printed fiber-reinforced polymers? What factors affect the magnitude of the porosity? By controlling those factors, can the porosity be reduced or eliminated?585Inter-Bead PorosityInternal PorosityFigure 1, Sample extruded on the BAAM at ORNL's MDF1.1 BackgroundThe large scale AM systems studied at the MDF are single screw material extruders. Because this technology was developed and modified from equipment used in the Injection Molding (IM) industry, a first step to understanding the issues facing AM is to examine similar issues within IM. A reference book on the subject, “Polymer Extrusion”, [1] dedicated several sections to the issue of air entrapment causing voids in molded parts. The text describes four methods of troubleshooting or reducing the voids.1. Adjust the temperature of the barrel and/or screw. 2. Increase the die pressure to keep the gases in solution. 3.

The text describes four methods of troubleshooting or reducing the voids.1. Adjust the temperature of the barrel and/or screw. 2. Increase the die pressure to keep the gases in solution. 3. Change the geometry of the screw or barrel to more rapidly pressurize the melting polymer. 4. Design a vent to extract gases prior to the final extrusion.From these steps we can discern three key factors that carry over to AM: trapped gasses, temperature, and pressure. Unfortunately, in AM polymers are extruded from the printhead directly into atmospheric pressures rather than into a pressurized mold like in IM. This means that even if the internal pressure was dramatically increased, the material would still be molten when exposed to atmospheric pressure allowing trapped gasses to leave solution and expand. Temperature and gas entrapment are readily controllable as can be seen in the field of polymer foams.

Temperature and gas entrapment are readily controllable as can be seen in the field of polymer foams. For example, Nikitine wrote in 2009 [2] how increasing gas solubility and temperature increased porosity. While the purpose of this study is to reduce porosity, not magnify it, papers such as this are a valuable resource into understanding one method through which porosity can form.Points three and four require redesigns of the extruder screw and barrel but are still worth considering. “Polymer Extrusion” provides an overview of both modified and unmodified single screws. A basic screw consists of three regions along the length of the screw: feed, compression,586and metering (Figure 2). Adding a vent to allow gasses to escape the system requires adding a low pressure region after the initial compression region, but before the final metering section.

Adding a vent to allow gasses to escape the system requires adding a low pressure region after the initial compression region, but before the final metering section. Figure 3 shows two methods for extracting those gasses from the venting section once the material is depressurized and the gasses have been allowed to leave solution. One option is a simple port on the side of the barrel. This is the simplest solution but requires the modification of the barrel itself as well as a method to seal the vent port if the vent is to be disabled. The second is a hollow screw with a bore hole into the root of the screw channel allowing gasses to escape through the rear of the screw.

The second is a hollow screw with a bore hole into the root of the screw channel allowing gasses to escape through the rear of the screw. This method could potentially be retrofitted to a system with minimal changes to the barrel or other components assuming the original barrel was long enough.Figure 2, Basics of an Extruder Screw [1]Figure 3, Dual Vented Screw [1]As a further note on trapped gasses, research at Purdue [3] found that entrapped air is only a partial cause of the porosity seen within AM. After applying a vent porosity was reduced but not entirely eliminated. Non vented samples on this system ranged from 16-27 % porosity and the paper suggested high extrusion rates as another potential source of the porosity for the fiber reinforced polymers being tested.

Non vented samples on this system ranged from 16-27 % porosity and the paper suggested high extrusion rates as another potential source of the porosity for the fiber reinforced polymers being tested. This was attributed to high shear rates damaging fibers and not allowing trapped gases to escape the melt.1.2 Mechanical PropertiesIt was generally accepted within literature that porosity would reduce mechanical properties. For example, a 2000 paper on glass fiber thermoplastics in compression molding [4] described a series of tests to show how increasing molding pressure reduced porosity to around 1.9 % and described this as a still significant amount.

The previously mentioned paper from Purdue [3] mentions how porosity affects mechanical properties, and attributes the majority of the problem to the porosity between layers rather than to the porosity within the beads themselves.587A few papers that isolate internal porosities from the interlayer porosities were from related fields. One was a study using finite element analysis to study the effects of porosities on crack growth in polymer composites [5]. This paper showed that voids reduced stiffness and that under very specific types of crack growth could improve the tensile properties by increasing the strain- to-failure. A second paper [6] also focusing on FEA analysis showed a near linear relationship between increasing porosity and decreasing relative stiffness in polymer foams.The third was an experimental study looking at porosities within carbon fiber reinforced polymer laminates [7] and was specifically looking at porosities ranging from 0.8-3.4%.

In their study, all mechanical properties were lowered as porosity increased by between 0.9-20.8%, but shear was affected far more than tensile stresses. A fourth paper [8] on porosity within vacuum- bag fiber-reinforced polymer panels showed that between the porosities of 1.5% to 3.4% fracture toughness decreased by 13% and fatigue life dropped by 50%.1.3 Tested Systems and Parameters of InterestThe purpose of this study was to compare the porosity found within samples produced on five different single-screw extruder systems. These systems were the Big Area Additive Manufacturing (BAAM) system at the MDF, an IM machine and a separate vented screw test extruder both belonging to Techmer PM, a proprietary commercial system, and finally a few archived samples from an older selectively vented extruder. All samples analyzed for this study were ABS 20 % carbon fiber with the exception of the archived samples whose precise composition was not available.

All samples analyzed for this study were ABS 20 % carbon fiber with the exception of the archived samples whose precise composition was not available. Various extrusion parameters were controlled within each system allowing for a comparison of their effects on porosity. These parameters were drying methods, pressure, venting, and extrusion rates including the effects of changing extrusion rates in the form of transient vs steady state flow. The primary focus of the study was on the samples from the BAAM system. The other four extruders were each examined for the purpose of comparison or to study the affects of factors BAAM does not currently have the capability to adjust.1.4 Sample Descriptions by System1.4.1 BAAMThe BAAM is a large-scale single screw extruder produced by Cincinnati and housed at the MDF of ORNL.

The seven single bead samples produced for this study were approximately 1m (3ft) long and printed at varying screw speeds ranging from 50 rpm – 350 rpm in 50 rpm increments. These were then sectioned in three locations the first inch of the bead, the middle middle, and the inch at the end of each bead just prior to a large deformation resulting from the extruder head halting and depositing extra material as it lifted off the platform. This resulted in 21 sets of two samples: both surfaces exposed by sectioning made up each set. The three locations were chosen to examine the effects of transient flow at three times within the print run: as the polymer first began to flow, once the polymer reached steady state, and as the system slowed down at the end of the bead. The pellet feedstock for these prints was dried at approximately 65ºC (150ºF) for 12hr in a desiccant dryer.5881.4.2 Techmer PMTwo systems at the Clinton Plant of Techmer were used.

The pellet feedstock for these prints was dried at approximately 65ºC (150ºF) for 12hr in a desiccant dryer.5881.4.2 Techmer PMTwo systems at the Clinton Plant of Techmer were used. The first was a single screw injection molding machine, and the other was a similarly proportioned single screw extruder outfitted with a vent in the side of the barrel. One standard tensile bar was produced via injection molding with a final molding pressure of 13.5k psi. The screw feedrate was 13.7 m/min (9 in/sec). The sample was sectioned from the center of the tensile bar. Just prior to molding the bar, an “airshot” was performed where the polymer was ejected directly into atmospheric pressure from the extruder for a few seconds. A sample of the 3mm bead produced was collected. A sample from a second airshot performed on a vented test extruder was also collected.

A sample of the 3mm bead produced was collected. A sample from a second airshot performed on a vented test extruder was also collected. The vent was 0.41m (16in) from the output nozzle of the extruder and was open to atmospheric pressure (as opposed to an induced vacuum). Four beads from each airshot were mounted and analyzed. The pellet feedstock for these samples had been dried at 70ºC (158ºF) for 2hr without the use of desiccants.1.4.3 Commercial SystemThe samples from this commercial single screw extrusion system were provided for this study under the conditions that the specifics of the system remain unpublished for now. The three provides samples were multi-bead walls extruded from a 4mm nozzle at feedrates of 2k, 6k, and 10k mm/min (1.31, 3.94, and 6.56 in/sec respectively) were provided. Three beads at each feedrate were sectioned and analyzed.

Three beads at each feedrate were sectioned and analyzed. Similarly to the BAAM samples, both surfaces exposed during sectioning were mounted providing a total of 6 samples per feedrate.1.4.4 Archived SamplesTwo images from an older extruder were analyzed. The system that produced them had been capable of switching between vented and non-vented extrusion. The original print parameters of the samples were largely unavailable. What was known was that the samples were identical except one sample was vented, and the other was not.2. EXPERIMENTATION2.1 Sample PreparationEach physical sample from each system differed significantly in size and shape. The BAAM samples were single flattened beads measuring about 1.5cm wide by 2-3mm thick (Figure 4). The Techmer airshots from the two extruders were 3mm diameter cylindrical beads and the injection molded bar was a 4mm by 10mm rectangle. The commercial system provided multi bead wide but single bead tall walls.

The commercial system provided multi bead wide but single bead tall walls. Each sample regardless of dimensions was sectioned to approximately 1cm in length on an Allied TechCut 5TM Precision High Speed Saw using a diamond wafering blade at 1000rpm for the blade and a feedrate of between 0.021 mm/s (0.05 in/min) and 0.064 m/sec (0.15 in/min). A slower feedrate protects the internal fibers from shattering. If fibers are too badly damaged this can result in higher fiber pullout as fragments of broken fibers fall out of the the polishing procedure. These polymer matrix duringFigure 4, Two adjacent BAAM Samples mounted in a 32 mm epoxy puck589fragments can result in polishing defects because they can be drug across the surface during the polishing resulting in scratches of varying severity. Fiber pullout is also a cause of false positives during porosity measurement as the fibers leave behind small holes in the matrix.

Fiber pullout is also a cause of false positives during porosity measurement as the fibers leave behind small holes in the matrix. To determine if fiber pullout was a concern, a typical multi-bead wall of fiber reinforced ABS from the BAAM was sectioned in two distinct ways. One sample was sectioned at 0.021 mm/s on the TechCut, and the other was the was sectioned with a bandsaw. Both samples were mounted in epoxy, and the sample surface of each was ground and polished. The grinding worked from 250grit for 2min all the way to 4000k grit for 15min and removed approximately 3.2 mm (1/8th in) of material. They were polished with a 5µm diamond slurry for 25min. After both samples were imaged and analyzed (process explained next), their average porosity across multiple sections of each sample, each contained three printed beads, was found to be 17.5 % and 17.7 % differing by only 0.2 %.

It should be noted that no significant scratching occurred from fiber fragments, so the only concern was the effect of false positives from fiber pullout. This indicated that the grinding and polishing procedure removed enough material to minimize the negative effects of rough or rapid sectioning on the surface of the sample. An example of a pair of mounted and polished BAAM samples can be seen in Figure 4.2.2 Porosity Measurement through Optical Micrographs and Image AnalysisTwo optical microscopes were used to image the samples for this study; a Zeiss Axio Imager and a Leica DM4000M. At total magnifications between 200x-300x, the Axio provided clearer images with higher feature contrast that were easier to analyze compared to those provided by the Leica, but both were acceptable. The key controls used for the analysis were exposure time, lighting intensity, gamma correction, and multi-focus stitching. There was no universal set of settings that would work for all samples.

The key controls used for the analysis were exposure time, lighting intensity, gamma correction, and multi-focus stitching. There was no universal set of settings that would work for all samples. The goal during the image capturing process was to achieve as high a pixel-value contrast between porosity and all other features as possible such that a pixel value threshold could be detected that would isolate porosity within the image. Exposure time and lighting were set by starting them at the default setting and then adjusting them to have longer duration and higher intensities until the background polymer matrix was as close to the same pixel value as the fibers within the matrix. One concern during this adjustment was ensuring that the porosity was not being washed out.

One concern during this adjustment was ensuring that the porosity was not being washed out. This would occur depending on the average depth or size of the pores within the specific sample; shallow pores were easy to wash out and lose, so for some samples the optimal contrast between porosity and background could not be achieved. Gamma correction refers to a post-processing feature built into both microscopes used for this study. It allows the user to skew the distribution of pixel values captured by the microscope. A gamma value between 0 and 1 skews the resulting image to have far more pixel values assigned to the dark regions of the image (low pixel values), while a gamma above 1 results in an image where more of the pixel value range is assigned to the bright regions of the image. For the purposes of this study, the feature of interest was porosity which appears as shadows or regions of very low brightness/pixel values.

For the purposes of this study, the feature of interest was porosity which appears as shadows or regions of very low brightness/pixel values. Therefore, a gamma value of between 0.2 and 0.3 was used to further increase the contrast difference between porosity and all other features. Figure 5 shows an image captured on the Axio imager that clearly displays the desired high contrast between the dark regions of porosity and the bright background matrix and fibers (small white dots).590Figure 5. A small section of an image captured on the Axio ImagerThe Fiji distribution of ImageJ [9] was used to perform the image analysis. The base distribution would have been sufficient for the analysis that was completed. The following generalized process was used and adjusted for each set of sample images.1. Start by cropping a rectangular region of interest out of your total image that avoids polishing defects or other unwanted features.a.

Start by cropping a rectangular region of interest out of your total image that avoids polishing defects or other unwanted features.a. Rectangular sections do result in missing porosities near heavily curved surfaces but can be analyzed far more efficiently by converting these steps into a repeatable macro with Fiji’s built in macro creator.2. Set the image to be an 8-bit black and white image.a. Several of the following tools can only function on an 8-bit image3. Create a copy of the new 8-bit image. 4. Subtract a value of 1 from all pixels in the copy. This will result in all pixels with the value of 0 and 1 having the value of 0 and no pixels having the maximum value of 255.5. Return to the original 8-bit image and run the process “subtract background” with a radius of 5 pixels.a. This is intended to help smooth out any uneven lighting in the background of theimage that may negatively effect the thresholding of the image.6. Apply an automatic threshold.a.

This is intended to help smooth out any uneven lighting in the background of theimage that may negatively effect the thresholding of the image.6. Apply an automatic threshold.a. There are multiple options available and depending on the precise lighting and contrast you managed to achieve through microscopy, different ones will prove to be more effective.b. In this study images captured on the Axio used the default threshold while those taken with the Leica needed to use the “MaxEntropy” setting to accurately identify the porosity.7. This produces a binary image with all porosity being assigned a pixel value of 255 and the rest of the image being assigned a value of 0.8. Run the “Analyze Particles” tool. Chose the options for “show masks” and “include holes.” a. Optional: set a minimum particle pixel size to eliminate “salt and pepper” defects resulting from the edges of fibers or other minor surface defects that were included in the threshold.

Optional: set a minimum particle pixel size to eliminate “salt and pepper” defects resulting from the edges of fibers or other minor surface defects that were included in the threshold. This trades the error of overestimating porosity due to false positives for the error of not including porosities smaller than the cutoff.591b. For this study, for each set of images (with shared microscope settings), a minimum particle size was set by measuring size of several manually selected defects. This typically ranged from 100-150pixels.c. The option to “include holes” fills in pores that only appear as outlines after thresholding. This often would occur when the pores were very shallow or wide and the bottom of the pore would reflect back more light to the sensor, appear brighter than desired, and be missed by the threshold.9. Add the “mask” created in step 8 to the edited copy created in step 4.

Add the “mask” created in step 8 to the edited copy created in step 4. To determine the porosity as a percentage of surface area, divide the number of pixels with a value of 255 from the total pixel count of the image. a. Now you have a single image where all detected porosity has been assigned a pixel value of 255, and the remainder of the image remains intact (but reduced in pixel value by 1 compared to the original).Figures 6 and 7 show the before and after of these 9 steps.Figure 6. A rectangular section of a BAAM sample to be analyzedFigure 7. Porosity isolated as the only pixels at a value of 2553. RESULTS3.1 BAAMFigure 8 summarizes the collected data from the porosity analysis. The primary variables being examined were screw speeds (and therefor flowrates) and transient vs steady state effects.

The primary variables being examined were screw speeds (and therefor flowrates) and transient vs steady state effects. Seven screw speeds were tested and samples were taken from the first inch of the extruded material, the middle of the extruded bead, and just prior to a large material deformation formed as the extruder nozzle came to a halt at the end of each bead. The effects of transient flow can be clearly seen as the porosity at the start of each bead was consistently higher than further along the bead. A second pattern is that the porosity becomes significantly more erratic as the screw speed passes 200 rpm. As the screw speed ramps from 50 rpm to 200 the porosity decreases. It also decreases at the end of the bead relative to the start of the bead for this rpm range. After 200 rpm, the porosities increase again, but with a higher variance than before.

It also decreases at the end of the bead relative to the start of the bead for this rpm range. After 200 rpm, the porosities increase again, but with a higher variance than before. The data at the end of each bead was inconclusive5922 mm2 mmas to whether transient effects occur as the system comes to a halt. This could be because the samples were taken from just prior to the large geometric deformation of the bead caused by the print head halting, lifting, and dragging material along with it. Any potential end effects could be caught up in that deformation but taking samples from within the deformation would result in measurements with far more unknowns. Overall, porosities ranged from 4.3 % (middle of the bead at 250 rpm) to 7.5 % (start of the bead at 50 rpm).BAAM: ABS 20%CFy t i s o r o P %8 7 6 5 4 3 2 1 0StartMiddleEnd50100200 250 Screw Speed (RPM)150300350Figure 8.

Porosity as a function of Screw Speed and Sample Location along the extruded bead3.2 TechmerThe Techmer samples were produced on injection molding equipment allowing a direct comparison to the industry from which pellet fed AM is derived. The low porosity molded bars show how gasses can be kept dissolved within the polymer by maintaining pressure, and the high porosity airshots show how much gas was contained within that polymer. All three samples were extruded from the same batch of pellet feedstock on the same day no more than 20min apart. This is important to note considering the variance in porosities produced across these samples. The molded bar (Figure 9) had only 0.1 % porosity, while the airshot produced from the same machine seconds prior averaged 16.7 %. The vented airshot averaged 0.3 % which is still higher than the molded bar but is 98% lower than the non-vented airshot.

The vented airshot averaged 0.3 % which is still higher than the molded bar but is 98% lower than the non-vented airshot. Figures 10 shows the high porosity within a non-vented airshot while Figure 11 shows a relatively pore free vented airshot. Table 1 summarizes the measured porosities. It should be noted that the porosity of the Techmer airshot was close to 17 % while the BAAM samples produced with the same material maxed out at about 7.5 %. A few possible explanations include moisture content, nozzle size, or extrusion feedrate.

A few possible explanations include moisture content, nozzle size, or extrusion feedrate. Of these, the most likely variable is moisture content because the BAAM samples were dried for 6 times the duration at a similar temperature with the aid of desiccants.Table 1, Collected data from Techmer samplesAverage % Min % Max %Molded Non-Vented Vented0.1 16.7 0.3n/a 15.5 0.1n/a 17.8 0.4593Figure 9, Techmer Molded BarFigure 10, Non-Vented Techmer AirshotFigure 11, Vented Techmer Airshot3.3 Commercial SystemThe primary purpose of the samples from the commercial system was as a second look at the effects of flow rates on porosity to compare to BAAM. Figure 12 shows a box and whisker plot of the porosity at each tested feedrate. As the feedrate increased, so did the porosity. At the highest feedrate of 10k mm/min there was also an increase in the variance of the porosity from 1.7% at 2k mm/min and 1.9% at 6k mm/min to 4.5%.

As the feedrate increased, so did the porosity. At the highest feedrate of 10k mm/min there was also an increase in the variance of the porosity from 1.7% at 2k mm/min and 1.9% at 6k mm/min to 4.5%. Figures 13-15 show how the porosity can be seen to increase as feed rate also increased. The bead depicted was chosen for being closest to the average porosity for that rate. Table 2 shows the average, minimum, and maximum porosities at each feedrate. The control scheme of BAAM is based on screw speeds while the commercial system controlled the feedrate of the material. Going forwards, further studies could include additional measurements to allow for a measurement of volumetric flowrate which would allow for a more direct comparison between systems.

Going forwards, further studies could include additional measurements to allow for a measurement of volumetric flowrate which would allow for a more direct comparison between systems. This would be particularly useful when trying to understand why the commercial system showed exclusively an increase in porosity as feedrate was increased, but BAAM showed two regions of differing behavior at low vs high feedrates.5943 mmFigure 12, Porosity % and Variance of a Commercial SystemFigure 13, Single bead at 2k mm/minFigure 14, Single Bead at 6k mm/minFigure 15, Single bead at 10k mm/minTable 2, Summarized Porosity data from Commercial SystemAverage % Min % Max %2k mm/min (1.31 in/sec)3.22.23.96k mm/min (3.94 in/sec)5.54.46.310k mm/min (6.56 in/sec)6.43.88.33.4 Archived Sample ImagesThe archived images from this system are valuable because they provide an example of the effectiveness of venting on porosity.

Table 3 lists the porosities for the two archived samples depicted in Figures 16 and 17. Using a vented system the overall porosity dropped from 18.4% to 6.6%, a 64% total reduction. These archived samples show that even after venting, porosity is not always eliminated as efficiently as it was in the Techmer samples. The reason for this discrepancy could not be determined due to the limited data on the archived samples.595Table 3, Porosity for each Archived SampleNon-Vented VentedPorosity % 18.4 6.6Figure 16, Archived Non vented sample: 18.4% PorosityFigure 17, Archived Vented Sample: 6.6% Porosity4. CONCLUSIONS4.1 FeedratesThere are clear trends between feedrates and porosity in both the commercial system and BAAM samples.

CONCLUSIONS4.1 FeedratesThere are clear trends between feedrates and porosity in both the commercial system and BAAM samples. In both cases, there was an increase in the average porosity and the variance of porosity within each set of samples at high feed rates, but BAAM suggests that this might not hold true at low feedrates.4.2 TransientsThe positional data from the BAAM samples showed that transient flow results in higherporosity. Further research needs to be done to determine the duration of these effects.4.3 Gasses and MoistureComparing the non-vented airshot from Techmer to the injection molded bar indicates that most of the porosity can be attributed to the pressure the samples were allowed to solidify at.

At the high pressures of the mold, any dissolved gasses or moisture appear to have remained in solution even after the solidified part was removed from the pressure chamber, while they expanded to form porosity as soon as they were exposed to atmospheric pressures during the airshot.4.4 VentingOf all the parameters tested, venting showed the most promise for reducing porosity with totalreductions of 64% in the archived samples and 98% for the vented Techmer airshots.5965. REFERENCES1. Rauwendaal, Chris. “Polymer Extrusion”. Carl Hanser Fachbuchverlag, 2014. 2. Clémence Nikitine, Elisabeth Rodier, Martial Sauceau, Jean‐Jacques Letourneau, Jacques Fages, “Controlling the structure of a porous polymer by coupling supercritical CO2 and single screw extrusion process” in Journal of Applied Polymer Science, Sep. 15, 2009, https://doi.org/10.1002/app.310313. DeNardo, Nicholas M., "Additive manufacturing of carbon fiber-reinforced thermoplastic 939. Open4. Greger Nilsson, S.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCharacterizing Material Transition for Functionally Graded Material Using Big Area Additive ManufacturingZeke Sudbury1,2, Chad Duty2,3, Vlastimil Kunc3, Vidya Kishore2, Christine Ajinjeru2, Jordan Failla2, John Lindahl31 Cincinnati Incorporated, Harrison, OH 2Mechanical, Aerospace, and Biomedical Engineering, University of Tennessee, Knoxville, TN 3Manufacturing Demonstration Facility, Oak Ridge National Laboratory, Oak Ridge, TNAbstractThis study examines functionally graded materials (FGM) on a polymer based large scale additive manufacturing system. FGM utilizes a less expensive material with sub-optimal mechanical properties for the majority of the part, and uses more expensive higher performance material in selected areas.

FGM utilizes a less expensive material with sub-optimal mechanical properties for the majority of the part, and uses more expensive higher performance material in selected areas. This process aims to optimize cost with weight and mechanical performance. FGM is already used a variety of industries, but is not common place in additive manufacturing, specifically large scale additive manufacturing like Cincinnati Incorporated’s Big Area Additive Manufacturing (BAAM). BAAM can use a variety of plastic injection molding and extrusion style polymer pellets, which allows it to use both commodity materials and high performance engineering polymers. This study is an initial assessment of FGM using glass fiber reinforced ABS and carbon fiber reinforced ABS, and characterizes the performance of a density gradient shape function to characterize the blending of materials.I.

BackgroundBig Area Additive Manufacturing (BAAM) is a large scale 3-D printing system that issignificantly larger and faster than most filament-based polymer extrusion systems. BAAM has a build area of approximately 2.5 meters by 6 meters with 2 meters of vertical travel. A high throughput, single screw extruder is mounted to the gantry that deposits beads of material ranging from 2.5 mm to 9 mm in diameter at a rate of 40 kg per hour. In addition, BAAM uses injection molding grade thermoplastic pellets instead of wound filaments. The most commonly used pellets are carbon fiber reinforced acrylonitrile butadiene styrene (CF-ABS), which range in price from $10 per kg to $20 per kg based on fiber loading and manufacturer. Glass fiber reinforced ABS (GF-ABS) has inferior mechanical and thermal properties, but costs around $2- $4 per kg. [1]738Figure 1.

Glass fiber reinforced ABS (GF-ABS) has inferior mechanical and thermal properties, but costs around $2- $4 per kg. [1]738Figure 1. Big Area Additive Manufacturing system front view [2]One particular benefit of additive manufacturing is the ability to tailor mechanical andthermal properties to a specific application. The local properties of a component can change based on internal structures, infill designs, and even material selection. This study is an initial examination of creating tailored properties using functionally graded materials with BAAM. Functionally Graded Materials (FGM) allows the use of a less expensive material with sub- optimal mechanical properties for the majority of the part, and a higher performance material in selected areas. This process aims to optimize cost with weight and mechanical performance.

This process aims to optimize cost with weight and mechanical performance. One basic example of an FGM application is a simply supported fiber-reinforced-composite beam that transitions from pure fiberglass at one end, to pure carbon fiber at the center, and back to pure fiberglass at the other end. In this example, the desire is to minimize deflection at the center of the beam by replacing the cheaper glass fiber reinforced material with a stiffer carbon fiber reinforced material. [3]Figure 2. Illustration of a fiber-reinforced-composite simply supported beam that transitions from fiberglass to carbon fiber to fiberglass again, where white lines represent fiberglass and blacklines represent carbon fibers . [4]The same principle of a simply supported beam with FGM can be applied to BAAM tooptimize weight and cost of a printed part.

[4]The same principle of a simply supported beam with FGM can be applied to BAAM tooptimize weight and cost of a printed part. A beam made entirely of GF-ABS material is five times cheaper than a CF-ABS beam, but weighs more, is more prone to warping, and will deflect more under load. A FGM beam aims to optimize the cost, weight, thermal, and mechanical properties.The goal of this experiment is an initial assessment of the ability to transition materials using BAAM to create FGM structures. Since BAAM utilizes a discrete pelletized feedstock, it is possible to vary the composition of the extruded material by selectively mixing the pellets in739the feedstock section of the BAAM system.

In addition, BAAM uses a single screw extruder, and sudden change in the feedstock materials will mix in the extruder to produce an output that gradually transitions between the two end states In the case of this study, the feedstock material is changed from glass filled ABS to carbon filled ABS during the extrusion process, and a characteristic shape function was defined to depict the transition between the materials. The shape function establishes the lag time between the input of the pellets and the output of the extruder. The shape function illustrated in Figure 3 estimates the fiber content of the extruded polymer at any given time, with the first vertical dashed line representing the lag time, or time required for the change in material to start occurring. The second dashed line represents the saturation time, which is the time required to reach a steady state value following the transition.Figure 3. Characteristic shape function that defines the transition in material.II.

Characteristic shape function that defines the transition in material.II. Experimental ProcedureA. MaterialsThe pellet feedstock materials utilized for this study were 20% by weight Glass Fiber- ABS (GF-ABS) and 20% by weight Carbon Fiber-ABS (CF-ABS) were supplied by Techmer Engineered Solutions. The materials were dried at 65ºC for four hours.B. Extrusion systemThe Cincinnati Incorporated BAAM system utilizes a custom, single screw extruder withfive heating zones and a single feed hopper (Fig. 4). The barrel is approximately 24-inch-long with a screw specially designed for processing fiber filled ABS. The screw has a maximum speed of 400 RPM, which correlates to a mass throughput of ~40 kg per hour.For this experiment, pelletized GF-ABS material was directly added to the hopper andextruded at a screw speed of 100 RPM (10 kg/h).

At this speed, the extruder has a residence time of 24 seconds, which is a measure of the amount of time required for a pellet to travel through the extruder. A sampling rate of 8 seconds was selected to capture three data points during a given pellet’s residence time in the extruder. In order to minimize unintentional mixing of pellets of different composition within the hopper, the transition between materials occurred in a “starve feeding” condition, in which all of the pellets of a given composition were completely consumed within the hopper (partially exposing the extrusion screw intake region) before a charge of pellets of a different composition were introduced. This approach insured that740materials mixed in the extruder rather than inside the hopper. In this case, the starting material was GF-ABS with a red dye to visually distinguish it from the 500 mL charge of black CF-ABS.

In this case, the starting material was GF-ABS with a red dye to visually distinguish it from the 500 mL charge of black CF-ABS. Figure 5 illustrates the transition in the extruded material from completely red GF-ABS (far left), to a mixed GF/CF-ABS (middle), to a saturated (predominantly black) CF-ABS on the far right. The transition represented in Figure 5 occurred over a period of 80 seconds, with 8 second intervals between samples.Figure 4. BAAM extruder.Figure 5. Extruded samples transitioning from glass fiber filled ABS (right) to carbon fiber filled ABS (left).In order to quantify the transition between materials and generate a characteristic shapefunction, the fiber content in each of the samples was measured with Thermogravimetric Analysis (TGA). TGA measures the weight loss of a small sample of a material as it is heated. The samples in this study (Figures 5,6) were heated to 800ºC at a rate of 10ºC per minute in air.741Figure 6.

TGA measures the weight loss of a small sample of a material as it is heated. The samples in this study (Figures 5,6) were heated to 800ºC at a rate of 10ºC per minute in air.741Figure 6. (Right)Universal Analysis Q500 TGA system [4]. (Left) Samples in TGA pansIII. ResultsThe fiber content of each sample can be quantified through TGA by analyzing the weightloss characteristics of each sample. In each case, the ABS matrix completely burns off in two stages; first styrene acrylonitrile (SAN) around 350C, followed by polybutadiene (PB) around 550ºC. The remaining sample is the combined weight of the glass or carbon fibers, which should be roughly 20% based on the average feedstock designations. As the temperature increases above 600ºC, the carbon fibers begin to pyrolize in air, leaving the only the remnant ash and glass fibers behind. Figures 7 and 8 show TGA results for the GF-ABS and CF-ABS pelletized feedstock materials, respectively.

Figures 7 and 8 show TGA results for the GF-ABS and CF-ABS pelletized feedstock materials, respectively. As expected, the samples experienced around 80% weight loss after the ABS burned off at 550ºC. The glass fibers remained constant around 20.5% of the starting weight. The carbon fibers burned off past 550ºC, but the fibers left behind 4.725% weight in ash.Figure 7. TGA of GF-ABS pellets. ABS is burned off around 550ºC and the fibers remain in the tray742Figure 8. TGA of CF-ABS pellets. The carbon fibers burn off and leave residual ash at 800ºC.The TGA plots for the extruded samples shown in Figure 5 are collected in Figures 9 and 10 below. The sample number in the legend corresponds to the time stamp and respective residual weight % listed in Table 1.Figure 9a. TGA plots of sample 1-10 with values for residual fiber shown in the upper left hand743Figure 9b. Section view of plots showing the difference in weight loss percent for each sample.Table 1.

TGA plots of sample 1-10 with values for residual fiber shown in the upper left hand743Figure 9b. Section view of plots showing the difference in weight loss percent for each sample.Table 1. Weight percent values at 800C for each sampleSample GF pellet 1 2 3 4 5 6 7 8 9 10 CF pelletTime (s) Weight %N/A 8 16 24 32 40 48 56 64 72 80 N/A22.5 22.42 22.09 22.14 22.5 18.21 13.35 10.38 10.36 8.518 9.928 4.73Since the remaining weight percent value at 800ºC (Table 1) correlates to the amount of glass fiber and residual ash for a particular point in sampling time, a system of equations was developed to relate the weight loss percent in Table 1 to fiber concentration.𝑋𝑋 +4.73 22.50𝑌𝑌 = 𝑊𝑊𝑋𝑋 + 𝑌𝑌 = 22.50 744X = % GFY = % CFW = weight percentEquation 1. System of equations to correlate weight percent values in Table 1 to glass and carbon fiber distributions in a sample.

System of equations to correlate weight percent values in Table 1 to glass and carbon fiber distributions in a sample. In the equation, X and Y are the unknown values of glass and carbon fiber percent, respectively; W is the value obtained from TGA.The difference between the GF-ABS baseline and weight percent value of a samplecorrelates to the amount of glass fiber remaining after the carbon fibers have burned off. However, because carbon fibers leave residual ash while burning off, the difference is corrected by a linear factor of the carbon fiber pellet baseline of 4.73. Table 2 lists the glass and carbon fiber concentrations in each of the samples as well as the initial feedstock materials. The glass fiber concentration was plotted in Figure 10 against time to observe time dependency of fiber concentration (i.e. the shape function illustrated in Figure 3).

The glass fiber concentration was plotted in Figure 10 against time to observe time dependency of fiber concentration (i.e. the shape function illustrated in Figure 3). Note that as the percentage of glass fiber content drops, the percentage of carbon fiber content increases accordingly.Table 2. Weight percent values at 800C for each sample and associated fiber concentrationsSample Time (s) Weight % X (GF) Y (CF) 22.5 GF pellet 22.42 1 22.09 2 22.14 3 22.5 4 18.21 5 13.35 6 10.38 7 10.36 8 8.518 9 9.928 10 4.73 CF pelletN/A 8 16 24 32 40 48 56 64 72 80 N/A22.50 22.40 21.98 22.04 22.50 17.07 10.92 7.16 7.13 4.80 6.59 0.010.00 0.10 0.52 0.46 0.00 5.43 11.58 15.34 15.37 17.70 15.91 22.49745Fiber percent vs Time25.0020.00)%(t n e c r e p r e b Fi15.0010.005.000.000102030405060708090Time (s)Figure 10.

Fiber content vs Time plot using values from Table 2The shape function in Figure 10 indicates that the transition from glass fiber to carbonfiber reinforcement occurs roughly 32 seconds after the CF-ABS was introduced into the extruder. Also, the time required to reach a steady state of carbon fiber content is approximately 53 seconds. At a spindle speed of 100 RPM, the BAAM extruder travels at ~3 inches per second. Under these conditions, the BAAM system would deposit up to 8 feet of material during the time required to initiate a material transition (32 s). Once the material transition is initiated, an additional 5 feet of deposited material would be required to achieve a steady state condition for the second material.IV. ConclusionsThe unique pelletized feedstock of BAAM allows materials to be mixed during theprinting process, which can be used to create functionally graded materials.

ConclusionsThe unique pelletized feedstock of BAAM allows materials to be mixed during theprinting process, which can be used to create functionally graded materials. The study was an initial assessment on transitioning from glass fiber filled ABS to carbon fiber filled ABS. Glass Fiber-ABS was extruded, and then 500 mL of Carbon Fiber-ABS was introduced into the extruder. Samples were taken every 8 seconds after the introduction of carbon filled material. Thermogravimetric analysis was performed on the samples, and the weight loss at 800ºC was correlated to the glass/carbon fiber content of that sample. The glass fiber content was plotted against time to observe the characteristic shape function for the transition and to quantify the lag time before the transition was initiated (32 s) and the saturation time required to complete the transition (53 s).

Under the deposition conditions of BAAM, these times correspond to deposited lengths of 8 feet and 13 feet, respectively, to initiate and complete a material transition. These initial results have not been statistically verified, and could vary as tests are repeated. Future studies will focus on reducing both time constants by testing at higher RPMS, different input functions that pre-blend pellets, and multiple transitions .746V. Acknowledgements Research sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle, LLC.References [1] C. Duty, V. Kunc, B. Compton, B. Post, D. Erdman, R. Smith, R. Lind, P. Lloyd, L. Love. “Structure and Mechanical Behavior or Big Area Additive Manufacturing (BAAM) Materials.” Rapid Prototyping Journal, 23 (2) (accepted February 2016) [2] Digital image. Cincinnati Incorporated. Web. 14 July 2016.

aracterizing the Bulk & Flow Behaviour of LS Polymer PowdersStefan Ziegelmeierab, Frank Wölleckea, Christopher Tuckb, Ruth Goodridgeb, Richard HaguebaBMW Group, Rapid Technologies Center, Germany, Knorrstraße 147, 80788 MunichbUniversity of Nottingham, Mechanical, Materials and Manufacturing Engineering, Nottingham, UK, NG7 2RD Accepted August 16th 2013AbstractThe properties of laser sintering (LS) powders affect processability and the quality of parts manufactured. This study compared three different methods used to quantify both the static and dynamic powder properties – a Revolution Powder Analyzer, FT-4 Powder Rheometer and Hausner Ratio. The aim of the work was to identify the most reliable method to characterize powder properties in correlation to the dynamic conditions that occur during LS. The experiments focused on different particle size distributions of a cryogenically ground polyurethane powder compared to a standard polyamide 12 LS material, PA2200.

The experiments focused on different particle size distributions of a cryogenically ground polyurethane powder compared to a standard polyamide 12 LS material, PA2200. The results have led to a deeper understanding regarding powder interactions and therefore serve as input for material design and quality assurance.Keywords: Powder Bed Fusion, Polyurethane, Powders, Laser Sintering, Flow Properties, Bulk Behaviour.1. IntroductionThe term Additive Manufacturing (AM) covers a variety of technologies used to build parts directly from a three-dimensional digital design data. Powder bed fusion (PBF) of polymers also known as Laser Sintering is one of the main additive processes that can be used to produce functional parts for testing or low series production [1]. However, the performance of components, mainly built from Polyamide 12, is limited concerning their near series properties.

However, the performance of components, mainly built from Polyamide 12, is limited concerning their near series properties. Since there are many different requirements for polymer consumer products, it is clear that the current repertoire of available polymers in PBF is no longer sufficient. Thus one main aim in AM research is to gain deeper understanding of the requirements of new laser sintering polymers. The most common polymers used in PBF and therefore the focus for the majority of current research, are semi-crystalline polymers predominantly polyamides (PA 12 or PA 11) or polyamide-based compounds as well as some filled varieties [2, 3]. To broaden the field of application, the research presented in this work introduces a thermoplastic polyurethane powder, a type of polymer which has not been well reported in AM literature up to now.

To broaden the field of application, the research presented in this work introduces a thermoplastic polyurethane powder, a type of polymer which has not been well reported in AM literature up to now. In addition, the majority of methods to characterize PBF materials that are described were originally established for conventional manufacturing processes, for example injection moulding or the storage of powders in hoppers [4, 5]. The complex interactions over the whole process chain in PBF demand a deeper understanding of the requirements for various powder qualities, the interactions of process parameters between the non-intrinsic (particle shape and size, flowability), as well as the intrinsic (melting point, melt flow) properties of the polymers used.

As PBF powders are stored statically in a reservoir before being applied ina dynamic manner, it is important to understand the particle size distribution (PSD) and shape coupled with its ability to flow under the roll spreading conditions present during PBF. Most researchers report on particle shape or morphology, size distribution and tap or bulk densities of the powders investigated, but do not take the dynamic flow properties into account [6, 7]. In addition the methods used to quantify the packing efficiency of the PBF powders are highly dependent on the user and do not necessarily lead to reliable data. In order to set up criteria for quality assurance it is essential to have analysis methods in place that work independently from its user. Amado et al.

In order to set up criteria for quality assurance it is essential to have analysis methods in place that work independently from its user. Amado et al. introduced a powder characterization method and found that its dynamic quantification of PBF powder properties lead to a more sensitive differentiation in characterization in comparing different commercially available powders [8]. In correlating a standard PA2200 with a cryogenically ground thermoplastic elastomer this study also takes the effects of different shapes and PSDs into account.2. ObjectiveDue to the lack of quantitative information on the dynamic properties of PBF powders, the development of a new polymer powder coincides with high efforts in trial and error cycles. The aim of this paper therefore is to compare three different powder characterization methods giving information on the packing as well as dynamic behaviour of different PBF powders and to identify the most reliable.

In complementing the existing knowledge on powder characterization approaches it is aimed to find methods that help to reduce the effort in qualifying new powders for PBF.3. Experimental approach and methodsThe experiments focused on different PSDs of a cryogenically ground polyurethane powder compared to a standard polyamide 12 PBF material, namely PA2200.The different fractions for the ground powders were achieved within an air classification process leading to a sharp separation between the size distributions. F0 therefore refers to the standard distribution out of the manufacturing process, F25 and F45 respectively indicates that no fine particles below 25 µm or 45 µm are enclosed. Commercially available powders showing good flowability consist of spherical particles with a narrow distribution in particle size around D50=50-60 µm and a volumetrically low amount of fine particles at around 10 µm [9, 10].

PA2200 therefore serves as the benchmark material.Size distribution and particle shape were initially analysed, as described in 3.1. The three different methods used to quantify static and dynamic powder properties were a FT-4 Powder Rheometer (3.2), a Revolution Powder Analyzer (3.3), and Hausner Ratio (3.4). In order to obtain comparable results, all the experiments were completed in the same environmental conditions with an ambient temperature of 22°C and a relative humidity of 50 %.

In order to obtain comparable results, all the experiments were completed in the same environmental conditions with an ambient temperature of 22°C and a relative humidity of 50 %. The goal of this study is to compare these methods with respect to 4 different process steps during laser sintering before the exposure of the powder bed as follows below:Cohesion of powder affecting packing and flow efficiency o BFE (FT-4) o Avalanche Angle (RPA) o Δ Density ( ) (HR)Flowability of powder during powder layer application o SE (FT-4) o Avalanche Angle (RPA) o HRPacking efficiency of the powder inside the feeders and build chamber o CBD (FT-4) o Density (RPA) o Bulk (poured) Density (HR)The cohesion between the powder particles plays a significant role on the packing efficiency within the build chamber and the flow efficiency of powders during powder layer application [3, 11-13].

Previous studies which analyze powder properties show that cohesion and flowability are two properties that are not clearly distinguishable and most of the times they are considered to describe the same characteristic or at least characteristics that are similar in terms of their definition [5, 14, 15]. However, the decision in this study was to present the results for cohesion separately from flowability characteristics to correlate them to a certain process step in PBF. Therefore, cohesion was considered to characterize more the powder behaviour within the static storage in the reservoir of the laser sintering system, while on the other hand, flowability relates to the materials dynamic characteristics during application of a powder layer.3.1 Size distribution and particle shapeTo identify the physical characteristics of the powder particles (size and shape), a Camsizer XT was employed.

Using a high resolution camera system consisting of two digital cameras, one camera detects the size and characteristics of fine particles, while the second camera is optimized for the analysis of large particles [16, 17]. The PSD was identified by the frequency distribution q3(x) [%/μm] based on volume. Sphericity and Width/Length Ratio were two shape factors chosen to describe the shape of the particles. For a sphere or a circle, the SPHT and the Width/Length Ratio shows a value = 1.3.2 FT-4 Powder RheometerAn FT-4 Powder Rheometer was used to identify the dynamic flow and bulk properties of the powders. The operating principle of the FT-4 standard Stability Test is based on a reproducible dynamic flow pattern. The test assesses the powder sample’s tendency to change its bulk and flow properties when it is forced to flow by an helix rotating through its bulk both downwards (anticlockwise) and upwards (clockwise) [18, 19].

The values obtained by the stability test describe how stable and flowable the powder is during the measurement. As the energies and densities measured are directly dependent on inter-particulate forces the effects of the powders’ shape and size can be detected [14, 19].3.2.1 FT-4 Characterization IndexesThe Basic Flowability Energy (BFE) [mJ] is the energy required to establish a compressive high stress flow pattern in a conditioned, precise volume of powder. The BFE is calculatedfrom the energy consumed by moving the blade downwards through the powder bulk from the top of the vessel to the bottom. The shape of the blade is such that its movement is highly compressive and because the powder is “confined” at the bottom of the vessel, the compressibility of the powder plays a major role in the BFE [18]. Cohesive powders consist of fine particles that include several air voids inside the bulk.

Cohesive powders consist of fine particles that include several air voids inside the bulk. During the downward movement of the blade, the particles are forced to flow at the blade face and cover the air gaps that exist in between the particles creating a relatively short and localized stress transmission zone. On the other hand, non-cohesive powders are generally packed more efficient and the air gaps inside the bulk are reduced. Therefore the stress transmission zone and as a result the flow zone is extended further inside the bulk raising the value of BFE due to friction and high inter-particulate contact stresses. In this case, high flow efficiency is a result of high BFE for powders with different PSDs [14, 20, 21].The Specific Energy (SE) [mJ/g] is a measure of how powder will flow in an unconfined stress environment and defines the mechanical interlocking of the powder particles.

It is calculated in the same way as BFE but during the upwards movement of the blade, where the sample is not confined and can be lifted without restrictions. While powder compressibility can be very significant for the BFE, the SE is most dependent on the shear forces and inter- particulate mechanical interlocking. Due to the low stress environment, cohesion, particles shape and surface texture are the most influential properties [22].Conditioned Bulk Density (CBD) [g/ml] is the density of a precise volume of powder measured inside the vessel of the FT-4 after the initial conditioning cycle of the bulk. Bulk density is dependent on many physical properties, such as particle size and distribution, particle shape, particle surface texture and cohesive/adhesive forces.

Bulk density is dependent on many physical properties, such as particle size and distribution, particle shape, particle surface texture and cohesive/adhesive forces. As a result of these properties the particles are arranged inside the bulk [14, 19, 23].3.3 Revolution Powder Analyzer (RPA) – Flowability TestThe Revolution Powder Analyzer is a measurement device to detect the dynamic flow properties of powders, consisting of a rotating drum and a digital camera system. The rotating drum is manufactured from aluminum and has an inner diameter of 110 mm and 35mm width. It is connected by two glass sides allowing the system to capture images of the rotating bulk. The camera detects the surface and the area of the rotating bulk in predefined time steps [24].3.3.1 RPA Characterization IndexesAvalanche Angle [°] is defined by the maximum angle obtained on the free powder bulk surface before an avalanche occurs. The average value of all avalanche angles is then being calculated.

The average value of all avalanche angles is then being calculated. The RPA software calculates the flowability angle from the center point on the powder edge to the top of the powder edge. As a general rule it could be stated that, the higher the avalanche angle the poorer is the flowability and the more cohesive is the examined powder sample [8, 24].Volume Expansion Ratio (VER) defines the ratio between the expanded bulk volume (bulk density) inside the drum and the tapped bulk volume in the vessel, before filling the drum during the preparation stage (tap density). The preparation step consists of filling a 100 mlcylinder under manual tapping until the maximum powder compaction is achieved. The volume of the sample is exactly the same for every powder.

The preparation step consists of filling a 100 mlcylinder under manual tapping until the maximum powder compaction is achieved. The volume of the sample is exactly the same for every powder. The expanded volume is measured by image capturing as the sum of the area of every image pixel occupied by the powder multiplied by the width of the drum [8, 25].Density: The bulk density inside the drum is calculated by the mass of the powder sample divided by the expanded volume inside the drum [24].3.4 Hausner ratioThe Hausner Ratio (HR) is a well established method to characterize bulk and flow properties of powders [26, 27]. To determine the Hausner Ratio only two laboratory accessories are needed, a measuring cylinder and a powder funnel.

To determine the Hausner Ratio only two laboratory accessories are needed, a measuring cylinder and a powder funnel. The cylinder that was used for the HR measurements was 100 ml and the tapping of the powder was performed manually.3.4.1 Hausner Ratio Characterization IndexesThe HR describes the ratio of tapped and bulk densities and therefore provides a statement on the flowability of powders. It is calculated as follows [28, 29]:The Δ Densities ( ) describes the difference between the bulk and tap density and served to characterize the cohesiveness of powders within this study.4. Results & Discussion4.1 Powders Physical CharacteristicsTable 1, overleaf, provides an overview of the physical properties of the materials observed, including the particle size, shape, moisture content, the base polymer as well as their manufacturing processes.

These properties represent the most influential magnitudes on the PBF materials non-intrinsic properties.PA 2200 36.5 50.9 66.3 0.891 0.763 Polyamide 12 Manufacturing Process PrecipitationCharacteristic/Powder D10 [µm] D50 [µm] D90 [µm] Sphericity b/l Base MaterialTPU F0 TPU F25 TPU F45 47.7 31.1 17.9 62.8 52.2 45.7 84.2 76.2 74.1 0.888 0.879 0.855 0.775 0.77 0.762 Thermoplastic Polyurethane Cryogenic MillingTable 1: PBF powder samples characteristicsIn Figure 1 the SEM images for these materials can be seen to indicate the morphology and shape characteristics.TPU F0PA 2200TPU F25TPU F45Figure 1: SEM images (resolution: 100μm for all) of PA 2200 [11] and the 3 TPU fractions [BMW]4.2 Particles size and shape characteristicsApart from the moisture absorption, other characteristics which form the powders bulk and flow properties are the PSD and the particle shape characteristics.

As SPHT leads to a higher differentiation between the shapes of the powders observed, only the results for this value are shown. In Figures 3 and 4, the PSD and the shape characteristics (SPHT) plotted against the PSD are presented. Figure 3 represents the PSD vs. inner width xcmin of the particles.Particle size distribution43]m µ %/2[ 3 q10050100 xcmin [μm]150200TPU F0TPU F25TPU F45PA 2200Figure 3: PSDs of powder samples over inner widthThe PSD for TPU F0 is broader than it is for the other powders as a result of the milling process. After sifting the powders they align more and more to the aimed PSD of the benchmark material, the PA 2200. TPU F45 shows a D10–D50 span of 36.5 µm (D10 +23.48 % / D50 +18.95 % / D90 + 21.26 %) compared to PA 2200 with a span of 29.8 µm. The SEM images (in Figure 1) were used as input for differences in the surface morphology.

The SEM images (in Figure 1) were used as input for differences in the surface morphology. As a general remark, TPU fractions have edgy particles with non spherical shape, especially the fractions that include finer particles meaning F0 and F25, whereas PA 2200 particles are closer to an ellipsoid or circular shape with a smoother surface texture. The difference in shape can also be seen in the SPHT-Q3 graph (see Figure 4). The powder with the most spherical particles (SPHT<1 cf. Section 3.1) was the PA 2200 with an average SPHT = 0.891 followed by TPU F45 (SPHT = 0.888), TPU F25 (SPHT = 0.879) and TPU F0 (SPHT = 0.855) respectively, indicating that the fine particles represent the most irregular shaped.100Sphericity (SPHT)80]m µ %/60[ 3 Q4020000,2TPU F00,4 TPU F250,6 TPU F450,8 PA 22001Figure 4: Cumulative PSD over particles sphericity (SPHT)In addition, the SEM images support the results regarding the shape of the powder particles.

As a result of the above mentioned, it can be noted that the fractions of TPU with larger particles have rounder or more compact particles and that the precipitated PA 2200 generates more spherical particles with a smoother surface, while cryogenic milling creates particles with a jagged edged shape.4.3 Comparison of characterization methodsAccording to the objective introduced, this section provides a comparison of the results from the analysis with respect to the powders’ bulk and flow properties.4.3.1 Cohesion resultsFigure 5 summarizes the results from the cohesion measurements using the three different methods introduced.140,00) 4 - T F (J m, ) 0 1 / 1 ( * E F B)A P R(o, e l g n A e h c n a l a v A)R H(lm / g, 0 0 0 1 * ) s e i t i s n e D120,00100,0080,0060,0040,0020,00( Δ0,00PA 2200TPU F0TPU F25TPU F45BFE*(1/10), mJ (FT-4) Avalanche Angle, o (RPA) Δ (Densities)*1000, g/ml (HR)Figure 5: Comparison (materials and methods) of cohesion resultsIn order to interpret the results depicted in Figure 5, the important conclusion is that the results for the BFE, Avalanche Angle and the Δ Densities show the same trends in terms of cohesiveness.

The higher the Avalanche Angle and the Δ Densities (see Section 3.4.1) are the more cohesive a powder behaves and vice versa for the BFE [8, 15, 24, 30]. The results show that with a decrease in fine particles within the TPU fractions the cohesive character of the powder bulk decreases (BFE TPU F0 = 709.72 mJ / TPU F25 + 33.26 % / TPU F45 + 46.31 %). TPU F25 with its lower amount of fine particles already gets close to the cohesion characteristics of PA 2200. The BFE even indicates a lower cohesiveness. Here the index of BFE presents a misleading trend when comparing PA 2200 and TPU fractions. It is known from empirical data during processing the materials in the PBF systems that the TPUs act more cohesively than the PA 2200 material. The Avalanche Angle as well as the Δ Densities show a straight forward trend compared to the processing characteristics of the powders within the PBF process.

The Avalanche Angle as well as the Δ Densities show a straight forward trend compared to the processing characteristics of the powders within the PBF process. TPU F45 with the lowest amount of fine particles shows the same value for Δ Densities compared to PA 2200 and a 9.53 % higher Avalanche Angle indicatinga negligibly higher cohesiveness. This could be due to van-der-Waals interactions that represent the dominating adhesive forces at lower particle sizes up to 50 µm [5]. With a further increase in particle size the influence of these forces decrease due to the increasing influence of the weight forces of bigger particles [11, 31].4.3.2 Flowability resultsFlowability is the exact opposite value of cohesion. The more cohesive a material behaves the lower its flowability and vice versa. Figure 6 depicts the comparison of the flowability results again using the three methods introduced within the methodologies section.

Figure 6 depicts the comparison of the flowability results again using the three methods introduced within the methodologies section. Since it is known from literature and the results out of this study, that the VER does not help to distinguish between powders with different shapes and PSDs, the flowability results of the RPA refer to the Avalanche Angle as well [8]. This value shows the same trend as the VER with a lower standard deviation and therefore delivers a more significant differentiation between varying powders. Furthermore, in terms of cohesion, the higher the Avalanche Angle is, the higher the cohesiveness of the powder gets.

Furthermore, in terms of cohesion, the higher the Avalanche Angle is, the higher the cohesiveness of the powder gets. As a consequence a high Avalanche Angle results in poorer flowability.71,220)A P R61,2001,180(o51,160, ) 0 1 / 1 ( * e l g n A e h c n a l a v A) 4 - T F ( g / J m,E S43211,1401,1201,1001,0801,0601,04001,020PA 2200 SE, mJ/g (FT-4)TPU F0TPU F25 Avalanche Angle*(1/10), o (RPA)TPU F45HRFigure 6: Comparison (materials and methods) of flowability resultsConversely, the Specific Energy is an indication of the particles’ mechanical interlocking as it detects the amount of energy that the helix needs during its upwards movement in order to break the inter-particulate bonds of the powder bulk. The bulk is not confined on the top of the FT-4 vessel, so no compressibility forces are present.

The bulk is not confined on the top of the FT-4 vessel, so no compressibility forces are present. The higher the Specific Energy is the stronger is the mechanical interlocking and the inter-particulate bonds of the powder that result in a poorer flowability. Figure 6 depicts that Avalanche Angle and Specific Energy show the same trends within the TPU fractions. TPU F0 which has finer particles (F0 D10= 17.9 µm c.f. F45 D10= 47.7 µm) shows the highest values in Avalanche Angle and SE resulting in the poorest flowability of all the samples. This is due to the changes in the dominating adhesive forces within the powder bulk described in Section 4.3.1, as with an increase in particle size the weight forces lower the influence of van-der-Waals interactionsR Hand therefore allow the particles to slide along each other easier. TPU F25 shows higher values of Avalanche Angle and SE than PA 2200 and TPU F45, implying also poorer flowability in comparison.

TPU F25 shows higher values of Avalanche Angle and SE than PA 2200 and TPU F45, implying also poorer flowability in comparison. As PA 2200 and TPU F45 show similar PSDs (Section 4.2) the higher Avalanche Angle and SE of TPU can be attributed to the irregular shapes and rougher surface of its particles and the higher mechanical interlocking while flowing past each other.The HR results show a similar trend compared to the Avalanche Angle and SE results, with a higher standard deviation though the measurements where repeated three times to get a statistically resilient statement.

The deviation results out of the fact that the procedure is highly user dependent due to the manual tapping and consolidation phenomena between powder particles influence the repeatability.4.3.3 Packing efficiency results (Feeders and Build Chamber)Figure 7 shows the results from the density measurements indicating that for the TPU fractions a decrease in the fine particles leads to an increase in packing density.

Although it is known from literature that a certain amount of fine particles is needed for an efficient packing to fill in the gaps between the larger particles, this does not apply for the jagged edged shaped TPU particles [11, 14, 32].0,500,40lm / g0,30, y t i s n e D0,200,100,00PA 2200TPU F0TPU F25TPU F45CBD, g/ml (FT-4) Density, g/ml (RPA) Density, g/ml (HR)Figure 7: Comparison (materials and methods) of packing density resultsIn addition the positive effect of a certain amount of fine particles filling in the gaps between the larger ones might be hindered due to the dominating adhesive forces between them (i.e.

TPU F0) and may not distribute homogenously within the overall powder bulk (see Figure 8).Figure 8: Schematic view of effects on packing efficiency between cohesive and non cohesive powders following [32]The higher packing density of F45 (+ 9.8 % in RPA) compared to PA 2200 therefore results out of larger contact areas between the less spherical particles. RPA and HR almost show similar values for density and align more to the real density to be expected. This can be aligned to the fact that the bulk density in the FT-4 is measured after one conditioning cycle where the helix went through the bulk performing rearrangements in the bulk and particles collisions and therefore leads to a higher density in general. This effect is even more significant for the powders consisting of fine particles (i.e. TPU F0) as the CBD density reduces the adhesive effects within the powder bulk due to the external compressive force that is applied [13, 14].6.

TPU F0) as the CBD density reduces the adhesive effects within the powder bulk due to the external compressive force that is applied [13, 14].6. ConclusionsIn designing new powders for PBF the analysis methods introduced could bring added value in predicting the static as well as the dynamic powder properties. It was shown that it is important to also take the particle shape and size distribution into account, as an improvement in the bulk properties could be reached in modifying these properties especially for cryogenically ground materials. For TPU F25 an improvement of the flow efficiency as well as its packing behaviour was shown in sifting out a low amount of fines below 25 µm, while keeping material wastage low.Cohesion:The BFE (FT-4) is an accurate value to detect cohesion behaviour but showed some drawbacks when it is about comparing powders with significantly different shapes as well as PSDs.

Detecting cohesiveness the Avalanche Angle of the RPA leads to reliable and reproducible trends (low standard deviation) within this study. The Δ Densities (HR) did not lead to a clear differentiation between the cohesiveness of TPU F45 compared to PA 2200. The more the analyzed powders differ in their particle shape and surface morphology the less sensitive the differentiation between them might be.Flowability:In terms of quantifying flowability or mechanical interlocking and therefore an indication of the behaviour of a powder during the application of layer in PBF, the results show that forpowders with very fine particles as TPU F0, SE (FT-4) is an accurate value. However, in the comparison between TPU F25 and TPU F45 the SE values are very close and do not distinguish between powders with a similar shape sensitively enough within the materials observed. The RPA again showed reproducible and reliable results in detecting flowability.

The RPA again showed reproducible and reliable results in detecting flowability. The deviations of the HR results were higher than all the other methods applied. HR therefore is not the method of choice to obtain reproducible results.Packing efficiencyThe CBD (FT-4) indicates changes in density but does not necessarily lead to a statement on real bulk density and only serves as a method to compare changes attributed to varying PSDs (see TPU fractions). The Densities in the RPA and the HR showed low deviations and correlate more with the real density of the powder bulk within the feeder of a PBF system.7. References[1]LEVY, G., SCHINDEL, R., KRUTH, J.P.: RAPID MANUFACTURING AND RAPID TOOLING WITH LAYER MANUFACTURING (LM) TECHNOLOGIES, STATE OF THE ART AND FUTURE PERSPECTIVES. In: CIRP Annals - Manufacturing Technology 2 (2003), p.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCHARACTERIZING THE INFLUENCE OF PRINT PARAMETERS ON POROSITY AND RESULTING DENSITYJames Brackett1, Dakota Cauthen2, Justin Condon2,3, Tyler Smith2,3 Nidia Gallego3, Vlastamil Kunc2,3, Chad Duty2,31 Energy Science and Engineering, University of Tennessee – Knoxville, TN 2 Mechanical, Aerospace, and Biomedical Engineering, University of Tennessee – Knoxville, TN 3 Oak Ridge National Laboratory, TNAbstractExtrusion deposition additive manufacturing produces parts with inherent porosity, whichtypically manifests as easily accessible voids between beads. This open porosity can also be accompanied by voids within the beads themselves, and both types can impact a part’s desired performance.

This open porosity can also be accompanied by voids within the beads themselves, and both types can impact a part’s desired performance. Porosity is influenced by a variety of factors, including infill percentage, layer height, nozzle diameter, print speed, and raster orientation. While their influence on mechanical properties and porosity have been studied previously, there has been minimal work connecting print parameters to porosity and subsequently to mechanical performance. This study investigates the relationships between print parameters, volumetric porosity, and mechanical performance. In addition, this study measures both open and closed porosity through use of a helium pycnometer rather than image analysis of a cross-section.

In addition, this study measures both open and closed porosity through use of a helium pycnometer rather than image analysis of a cross-section. Thus, this study will identify correlations between the volumetric density of parts and the resulting mechanical performance as a function of print parameters.IntroductionExtrusion deposition Additive Manufacturing (AM) techniques vary in extrusion processand materials, but one of the most common processes is Fused Filament Fabrication (FFF) of thermoplastic polymers. While not limited to small-scale, the technique is most often associated with desktop printers found in homes and labs across the nation. The popularity of commercial units is largely due to FFF’s ease of use and affordability, but it still maintains important applications in modelling, prototyping, and small-batch production [1]. However, FFF does encounter issues with resolution and porosity, which the scale of the layer-based structure inherently introduces [2].

However, FFF does encounter issues with resolution and porosity, which the scale of the layer-based structure inherently introduces [2]. From a material standpoint, FFF is compatible with a wide range of thermoplastics, but one of the most popular choices is Acrylonitrile-Butadiene-Styrene (ABS). ABS material properties provide easy processing conditions for FFF while its price-point is lower than many engineering plastics, providing a useful middle-ground [3]. Thus, ABS is an ideal candidate for examining the FFF process and its inherent influences on the properties of printed structures.There have been numerous investigations into the influence of environmental conditions, material characteristics, and printing parameters on the FFF process and the resulting properties of printed structures. To demonstrate environmental dependence, Sun et al investigated the effects of variation in processing temperatures on the quality of bonding between adjacent rasters.

To demonstrate environmental dependence, Sun et al investigated the effects of variation in processing temperatures on the quality of bonding between adjacent rasters. Through measurement of neck formation and flexural loading, they demonstrated the766importance of controlling cooling to improve mechanical performance through increased raster bonding [4]. While ABS is one of the more popular choices for FFF systems, they are compatible with a variety of thermoplastics, each of which has its own distinct properties and suggested printing parameters. For example, PolyLactic Acid (PLA), another extremely common FFF material, can vary from one brand of PLA to the next. Wittbrodt and Pearce compared the tensile performance of five different colors of PLA at four different extruder temperatures, and they found that each color had a distinctly different performance with extruder temperature further altering tensile strength as a result of variation in crystallinity [5].

However, in a study of seven different thermoplastic filaments, it has been shown that the exhibited tensile strength is a direct product final mass as a function of the theoretical mass, regardless of material, which led the team to identify under-extrusion in the interior as the culprit for inconsistencies in tensile strength [6]. Even within a given set of environmental and material conditions, the chosen print parameters still further influence FFF printing. The number of layers within a part, for instance, was found to affect the mechanical performance such that observed moduli and strength values steadily increased until reaching a substantial number of layers (12 at 0.2 mm layer height) [7]. Similarly, Wu et al found a “sweet spot” in layer height experiments where a layer height of 0.3 mm significantly outperformed both a 0.2 mm and a 0.4 mm layer height [8].

Similarly, Wu et al found a “sweet spot” in layer height experiments where a layer height of 0.3 mm significantly outperformed both a 0.2 mm and a 0.4 mm layer height [8]. In a further exploration of Letcher’s work, Rankouhi et al demonstrated an improved performance at smaller layer heights and through optical microscopy, attributed the improvement to the reduction in the “air-gap to material ratio” that was seen at a 0.2 mm height vs the 0.4 mm height, meaning there was more material in the mesostructure of the 0.2 mm samples [9]. Each of these studies also incorporated a varying raster angle in their experimental process; unidirectionally, 0° raster angle outperformed a 45° and 90° orientation while Wu found the best performance in an alternating 0°/90° orientation compared to a 30°/-60° and 45°/-45° orientation [7-9].

An analytical study of more possible raster orientations found agreement with these results, showing a continual decrease in Ultimate Tensile Strength (UTS) with increasing raster orientation angle [10]. Another parameter of interest is the infill percentage. Alvarez et al studied the effectiveness of printing from 0% to 100% infills and found minimal performance improvements for infills from 50% - 98%, after which there was as significant jump [11]. Hossain et al investigated multiple build parameters (build orientation, raster orientation, raster width, contour width, and air gap between rasters), and while all were found to have significant impacts on the UTS, they also demonstrated that a negative air gap could be used to eliminate space between rasters for a nearly defect-free part with dramatically improved UTS [12].

While not a comprehensive list, there are obviously numerous parameters that can influence an FFF printed part and extensive study has been devoted to understanding each one’s contribution.767Although there have been numerous investigations into the effects of layer height, infill percentage, and porosity on mechanical properties, they generally consider only one or two of the three. In this work, we seek to examine all three simultaneous and investigate how each are connected. To that end, we examine the influence of layer height and infill percentage on the porosity of a printed sample before moving forward to consider the influence of the print parameters and resulting porosity on the UTS and elastic modulus of tensile specimens. This approach is intended to capture the effects of each parameter separately and the effect of their combined presence.ExperimentalPurple Matter Hacker Build Series ABS filament [16] was used to print all samples on a MakerGear M2.

The filament was received sealed and dried before being immediately used, and it was maintained in a dry environment during the printing process.Print ParametersAll print parameters other than those being studied were held constant. The extrusionmultiplier was set to 1.0, filament diameter of 1.75 mm, infill angle of (cid:3399)45°, bed temperature of 80°C, nozzle temperature of 235°C, print speed of 60 mm/s, solid infill speed of 20 mm/s, nozzle diameter 0.35 mm, and 0 perimeters. The parameters chosen for study were the infill percentage and layer height. Additionally, the chosen layer heights were obtained as ratios of nozzle diameter to enable a more direct comparison to other nozzle sizes.

Additionally, the chosen layer heights were obtained as ratios of nozzle diameter to enable a more direct comparison to other nozzle sizes. Table 1 shows the resulting layer heights and infill percentages used, each of which was printed with all four settings from the opposite row for a total of 16 parameter pairs.Table 1: Parameters Under InvestigationLayer Heights (mm)0.1550.1750.2000.235Infill Percentages85%90%95%100%Print and Sample GeometryFor this investigation, each parameter set was printed as a flat, rectangular plate. Thismethod was chosen to avoid the stress concentrations brought about by geometric discontinuities in the radial regions that are associated with printing dogbone tensile samples directly. Tensile samples were then machined from the plates using waterjet cutting in an orientation that preserved the (cid:3399)45° infill structure. Machining dimensions were taken from ASTM D638-14 Type V specifications [17].

Machining dimensions were taken from ASTM D638-14 Type V specifications [17]. The plates were 111.125 mm by 88.9 mm (4.375 in by 3.5 in) to allow for eight dogbones to be extracted and ensure statistical relevance. Samples that experienced damage or delamination, about one per plate, at any point were discarded. Furthermore, the chosen print parameters produced three plates that failed during printing due to768l l i f n i% 5 9e h td n at h g i e hr e y a lm m 5 7 1 . 0d n a m m 5 5 1 . 0h t iwl l i f n i% 0 0 1e h t: n o i t a n i m a l e dd e b. s e l p m a s d e t c a r t x e o te t a l p d e t n i r p m o r fs s e c o r p e h ts w o h s 1 e r u g i F. t h g i e h r e y a lm m 5 5 1 . 0 h t iwh t iw e t a l p e h t d n a, y a w h t a p g n i t t u ctt e j r e t a w e h t.

A testing rate of 1mm/min was used for a nominal strain rate of 0.1 mm/(mm*min). Extension was measured using an MTS LX 500 Laser Extensometer. Samples that fractured outside of the gage length were not included in data analysis.Results and DiscussionApparent and Volumetric Density MeasurementsFigure 2 shows the apparent density measured via helium pycnometry. The measured density for the filament is greater than that of the printed samples, but even with the distinct difference, the 0.155 mm layer height’s apparent density showed a less than 2% difference. Given that all other plates fall in between the two, the printing process appears to introduce minute amounts of closed porosity into the filament strands.

Given that all other plates fall in between the two, the printing process appears to introduce minute amounts of closed porosity into the filament strands. This is likely a result of both standard deviation in the measurement process and possibly a small amount of air infiltration within the melt inside the extruder.Average Density Measured via Pycnometry1.02501.02001.0150~ (')E (.) :§ 1.0100~ en ~ 1.0050 0 Q) ~ 1.0000 ~ <(0.99500.99000.985085% Infill- 0.155 LH85% Infill- 0.175 LH85% Infill- 0.200 LH85% Infill - 0.233 LHFilament Batch 1Figure 2: Apparent densities of four representative samples from printed plates and the filament used to print them.770Since FFF processes inherently produce structural porosity within a part, it was expected that helium pycnometry measurements would show a minimal difference and essentially measure the density of the beads within the piece.

While Figure 2 supports this, Figure 3 confirms that the “true” density of the printed dogbones was significantly lower than those obtained through pycnometry. For example, at the lowest point, the print density was only 82% of filament’s apparent density. Furthermore, the density of samples increased linearly with infill percentage. Figure 4 provides a closer picture of layer height’s more nuanced influence on density. Although the difference was minimal, 85% and 90% infills exhibited a gradual decrease in density with increasing layer height: a 3% decrease in the 8% infill set and a 5% decrease for 90% infill. This effect was not seen for the 95% and 100% infills, for they exhibited no correlation between layer height and density.Density vs Infill Percentage1.050Filament Density= 1.207 g/ cc------- , -----------------------e O.lSSmm Layer"'E 0.950 u -tlO -a.

Figure 5 exhibits a clear trend of increasing elastic modulus with increasing infill percentage, up to a maximum of 216% from the 0.233 mm layer height at 85% infill to the 0.233 mm layer height at 100% infill. However, it does experience a plateau effect moving from 95% to 100% infills. Once again, Figure 6 considers layer height’s influence on the modulus and yields two separate trends. As before, the 85% and 90% infills experience a decrease in modulus with increasing layer height – up to 47% and 40% respectively. However, no substantial difference is observed between layer heights at 9% and 100% infills. Figure 5 indicates that increasing infill past 95% yields diminishing returns on modulus improvement while Figure 6 shows that layer height remains important in determining the modulus until a threshold between 90% and 95% infill.772Young's Modulus vs Infill Percentage2.001.80 CU 1.60 0..

~ 15.00 =>10.00Iitt+ t+ •90% Infille 95% Infille 100% Infill5.000.000.1350.1550.175 0.195 0.215 Layer Height (mm)0.235Figure 8: The UTS of each print, grouped by infill percentage.Specific Moduli and UTSTo complete the investigation, the mechanical properties were considered as a function ofdensity. In Figures 9 and 10, the averages were plotted with both layer height and infill774percentage legends to establish relationships between both print parameters, porosity, and mechanical performance. For Figure 9, there is a relatively constant increase in specific modulus up to 192% at the 0.233 mm layer height at 95% infill. From here, even though the 100% infills exhibit a higher density, the lack of improvement in moduli results in a 6% decrease in specific moduli. Furthermore, a separate trend lies within each infill percentage.

Furthermore, a separate trend lies within each infill percentage. For the 85% and 90% infills, the specific modulus generally increases as layer height, and therefore porosity, decreases. However, for the 95% infill, the trend reverses and the increase in specific modulus is caused by an increasing density related to increasing layer height. As for the 100% infill, while there only two viable sample sets, they appear to cluster more tightly around a given value regardless of layer height. While similar trends are evident in Figure 10, they are not as distinct. Within both the 85% and 90% infill, specific UTS increases with decreasing porosity, but the 90% infill actually exhibits a lower specific UTS compared to the 85% infill. This is a result of significant densification in the 90% specimens without substantial UTS improvements. Both the 95% and 100% infills actually exhibit a performance improvement with decreasing layer height leading to an increased specific UTS.

Two legends were used, where color identifies infill percentage while shape of the point identifies the layer height.775Ultimate Tensile Strength vs Apparent Density35.0030.00e 85% Infill 1 1e .j. 1 e 90% Infill....... 25.00 ro ~ 20.00 ......... ~ 15.00 :J10.00e 95% Infille 100% Infill♦ 0.155 mm■ 0.175 mm0.200 mm5 .000.233 mm0.000.8000.8200.8400.8600.8800 .9000.9200.9400.960PApp (g/cm3)Figure 10: Comparison of specific UTS by plotting UTS vs density. Two legends were implemented, where color represents infill percentage while shape identifies layer height.Conclusions and Future WorkThe vast majority of porosity in printed parts is, as expected, seen to be open celled, asshown by comparing non-inclusive density measurements for the base filament and printed parts. The densest specimens exhibited 7% porosity compared to the filament used to print them.

The densest specimens exhibited 7% porosity compared to the filament used to print them. The most porous structures were approximately 19% porous, which demonstrates a 64% reduction in porosity from the most porous print parameter pair to the densest. Considering the significant porosity in all samples, simple measurements of mechanical performance using a filament’s base density are likely inaccurate and potentially misleading. As such, reporting porosity is critical to any analysis involving the effects of print parameters on mechanical properties.Increasing the infill density of a printed part reliably increases the density, elastic modulus, and UTS of it, although the elastic modulus does not benefit from increasing the infill percentage past 95%. Similarly, at high infill percentages, increasing the layer height of a print reliably produced an increase in density, modulus, and UTS.

Similarly, at high infill percentages, increasing the layer height of a print reliably produced an increase in density, modulus, and UTS. However, at the 85% and 90% infill settings, increasing the layer height increased the porosity of the print and hampered mechanical performance. Specific moduli and UTS demonstrated unique trends in that layer height controlled the mechanical performance at a given infill percentage. Ultimately, it indicates a direct influence of the print parameters on the mechanical properties of a part while also highlighting the influence of the porosity created by the parameters’ interactions.AcknowledgementsResearch sponsored by the U.S. Department of Energy, Office of Energy Efficiency andRenewable Energy, Industrial Technologies Program, under contract DE-AC05-00OR22725 with UT-Battelle, LLC.

Department of Energy, Office of Energy Efficiency andRenewable Energy, Industrial Technologies Program, under contract DE-AC05-00OR22725 with UT-Battelle, LLC. This work was supported in part by Oak Ridge Institute for Science and Education through the Higher Education Research Experiences Program (HERE).776References[1][2][3][4][5][6][7][8][9][10]B. Wendel, D. Rietzel, F. Kuhnlein, R. Feulner, G. Hulder, and E. Schmachtenberg, "Additive Processing of Polymers," Macromolecular Materials and Engineering, vol. 293, pp. 799-809, 2008 2008. S. C. Ligon, R. Liska, J. Stampfl, M. Gurr, and R. Mulhaupt, "Polymers for 3D Printing and Customized Additive Manufacturing," (in English), Chem. Rev., Review vol. 117, no. 15, pp. 10212-10290, Aug 2017, doi: 10.1021/acs.chemrev.7b00074. E.-K. Karahaliou and P. A.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCHARACTERIZING THE INTERNAL MORPHOLOGY OF TRANSITION REGIONS IN LARGE-SCALE EXTRUSION DEPOSITION ADDITIVE MANUFACTURINGJames Brackett1, Zaky Hussein2, Elijah Charles2, Tyler Smith3, Ahmed Hassen3, Seokpum Kim3, Vlastimil Kunc3, Chad Duty2,31 The Bredesen Center for Interdisciplinary Research and Graduate Education, University of Tennessee – Knoxville, TN 2 Mechanical, Aerospace, and Biomedical Engineering, University of Tennessee – Knoxville, TN 3 Oak Ridge National Laboratory, Oak Ridge, TNAbstractA dual-hopper feed system that was developed for the Big Area Additive Manufacturing (BAAM) system allows for transitioning between different materials while maintaining continuous deposition.

This technique creates a step-change in material feedstock by switching the pellet feeding system to alternate which hopper is currently supplying material, allowing for multi- material construction. The step-change in feedstock material produces a transition region that is characterized by a compositional gradient and blended internal morphology. Initial cross-sectional imaging of the transition region revealed a non-homogenous blend of materials with distinct domains of each material, likely due to incomplete mixing within the screw. This study used a carbon fiber reinforced acrylonitrile butadiene styrene (CF-ABS) and an unfilled ABS to characterize the internal structure and to correlate it to mechanical performance by tracking microhardness across cross-sections of the transition region.IntroductionLarge-Format Additive Manufacturing (LFAM) has seen a rapid increase in interest in recent years, evidenced by the increasing number of systems available.

Some examples include Cincinatti Incorporated’s Big Area Additive Manufacturing (BAAM) system featured in recent literature [1-3], Thermwood’s commercially available Large Scale Additive Manufacturing (LSAM) system [4-6], and even international efforts such as this custom prototype Polymeric Pellet-Based Additive Manufacturing (PPBAM) system [7].One of the more prevalent research areas is the introduction of multi-material capabilities. For example, recent developments with the BAAM have seen the introduction of a dual-hopper system capable of automated, on-the-fly material switching [2]. The system complements the pelletized feedstock by introducing a step-change in material feedstock then allowing the extrusion screw to blend pellets of differing composition and produce a continuous transition from Material A to Material B. Figure 1 shows an example schematic of the process, indicating key transition points (1A, 1B, 1C).

Figure 1 shows an example schematic of the process, indicating key transition points (1A, 1B, 1C). After purging Material A through the system to reach steady-state, the hopper switch would be initiated at 1A. From there, there is remaining Material A still in the extrusion barrel that must exit before the blended region begins at 1B. The transition continues until 1C where a steady-state Material B has been reached. This is also shown graphically with the y-axis indicating material composition and x-axis representing volume of extruded material.761Figure 1: A schematic representation of a printed material transition using the BAAM Dual-Hopper system where blue represents Material A and red represents Material B.This approach avoids some of the typical pitfalls of multi-material extrusion with thermoplastics by eliminating discrete material boundaries and the need to stop the extrusion process to switch materials.

However, the blended transition region presents other possible issues. As shown in Figure 2, cross-section images taken from the transition region where material blending occurs shows distinct regions of both Material A (acrylonitrile butadiene styrene, ABS) and Material B (20 wt % carbon-fiber filled ABS, CF-ABS). The heterogeneous internal morphology could have distinct effects on the mechanical performance of the transition region.Figure 2: The change in fiber content of a printed material transition from neat ABS to CF/ABS as a function of extruded volume. Images are cross-sections that were taken from the indicated point in the transition.Investigating the mechanical behavior of a single printed bead can be difficult, especially when trying to discern differences between regions in the cross-section. As such, hardness measurements provide a unique opportunity to investigate deformation resistance on a local scale using microhardness indentation.

As such, hardness measurements provide a unique opportunity to investigate deformation resistance on a local scale using microhardness indentation. Typically, thermoplastics are submitted to Shore hardness testing using a durometer [8], as was done by Vian and Denton in this study comparing the change in hardness with processing conditions [9]. For other studies, particularly those focused on microhardness, a Vickers Hardness system was used. Zheng and Ashcroft used microhardness indentation to investigate the effectiveness of various adhesives and their application to a part [10]. Turnbull and White utilized microhardness in an in-depth characterization study of unplasticized poly-vinyl chloride [11]. Another similar study saw a change in Vickers hardness with differences762in nanoparticle fillers using microhardness testing at a variety of applied loads [12].

Another similar study saw a change in Vickers hardness with differences762in nanoparticle fillers using microhardness testing at a variety of applied loads [12]. A comprehensive study of a dental composite material evaluated its resistance to deformation and established an increased stability in results with higher loads and longer dwell times [13].Characterizing the mechanical behavior of the transition region created when switching materials with the dual-hopper BAAM system is critical to understanding how the blended material composition will affect overall part performance. The first step in this process is understanding how the internal morphology changes with processing conditions and the effect of the regional morphology. Vickers microhardness indentation provides a tool well-suited for characterizing the limited area provided by a single-bead cross-section.

Vickers microhardness indentation provides a tool well-suited for characterizing the limited area provided by a single-bead cross-section. This study will utilize hardness measurements to represent local resistance to deformation and examine differences in performance based on regional composition. The effect of changes in processing conditions will also be investigated.Experimental MethodsAll samples were printed using the BAAM dual-hopper system. Three sample sets were created to isolate the effects of bot extrusion screw geometry and direction of material transition. Samples were extracted from locations of interest in each of the three transitions. Carbon fiber content was measured via Ultrasonic-Assisted Acid Digestion, as described in previous work [3, 14]. Vickers microhardness was measured at multiple locations on each cross-sectional sample.Print ConditionsPrint conditions and materials were constant for all printed samples.

Vickers microhardness was measured at multiple locations on each cross-sectional sample.Print ConditionsPrint conditions and materials were constant for all printed samples. Techmer Engineered Solutions provided both a neat ABS resin (HIFILL ABS 1512 3DP) and a 20 wt % CF/ABS (ELECTRAFIL ABS 1501 3DP) for sample production. Material feedstock was dried at 80 °C for at least four hours. Thermal processing conditions were kept constant regardless of which hopper was in use, maintaining a bed temperature of 100 °C and melt temperature of 250 °C. Using a 10 mm (0.4 in) diameter nozzle, each structure was printed with bead dimensions of 5 mm (0.2 in) by 14 mm (0.55 in).Sample PreparationAll material transitions were printed as a single continuous bead using a dual-hopper equipped BAAM at ORNL’s Manufacturing Demonstration Facility.

Three different parameter sets that varied either screw geometry or transition direction were used in this study: Non-mixing ABS to CF/ABS, Mixing ABS to CF/ABS, and Non-Mixing CF/ABS to ABS. All other system parameters were held constant. Cross-sectional samples 10 mm long were extracted using an Isomet 1000 Precision saw from five locations within each printed transition. The five cross- sections were chosen based on visual appearance and measured fiber content. One was taken from both the neat ABS and CF/ABS regions while three samples were taken from the transition regions. The specimens from the transition region were confirmed to be of similar fiber content to those taken from the other transitions by previous work [15].763Each of the fifteen cross-sections were embedded in an EpoxiCure 2 epoxy disc and polished with the same procedure.

A six-slot AutoMet 250 autopolisher was used to grind the reverse side of the discs flat using 240, 320, 400, 600 and 800 SiC grit paper in sequence. The cross-sectional surface was polished with 240, 320, 400, 600, 800, and 1200 SiC grit for 1 minute each with 2 lb central force. The sample holder was set to 60 RPM at a contrary rotation to the polishing platen (120 RPM). The surface was finished using a 6 µm diamond suspension with complimentary rotation of the sample holder for 5 minutes.Microhardness MeasurementsMicrohardness indentation was conducted using a Wilson VH1202 Micro Hardness Tester following ASTM E384 [16]. Based on literature [11, 13], a 100 gf load and 15 s dwell time was used for all indentations to reduce variance as much as possible. Vickers Hardness (HV) values were calculated by the instrument using the measured diagonals. The indentation pattern for each cross-section was a rough cross shape, as seen in Figure 3, with a 1 mm spacing.

The indentation pattern for each cross-section was a rough cross shape, as seen in Figure 3, with a 1 mm spacing. Each indent is circled in green. This pattern provided an unbiased selection of data points to analyze possible variance in the heterogenous morphology. The average HV and standard deviation were recorded for each cross-sectional sample. Any indentations hindered by porosity or sample defects were not considered. Average HV and standard deviations were plotted as a function of fiber content.Figure 3: An example of the pattern used when making microhardness indentations on the studied cross-sections. The indents are circled in green.764Optical MicroscopyEach cross-section was imaged with a Keyence VHX-5000 digital microscope. To capture a detailed image of the entire cross-section, the 2D-stitching software was used at 200x magnification. These images were used to assess whether an indentation had been hindered by sample defects.

These images were used to assess whether an indentation had been hindered by sample defects. Furthermore, the images were used to identify three distinct regions (neat ABS, mixed, and carbon-fiber rich) present in the cross-sections and determine in which region each indentation fell. Figure 4 provides an example of how each region was differentiated; blue, purple, and red represent the neat ABS, mixed, and fiber-rich regions, respectively.Figure 4: An example of the different regions observed in sample imaging. Color-coding indicates the region an indent fell in.Results and DiscussionRegional Variance of MicrohardnessThe measured HV of each indentation on a representative cross-section were compared to analyze the change in HV compared to change in surrounding material composition. Using the same sample from Figures 3 and 4, Figure 5 shows the HV of two indentations from each of the three regions.

Using the same sample from Figures 3 and 4, Figure 5 shows the HV of two indentations from each of the three regions. Once again, the blue, purple, and red signified neat ABS, mixed, and fiber-rich regions respectively. The results indicated a clear increase in HV with increasing fiber content in the surrounding areas. This trend was consistent across all cross-sections, with fiber-rich regions displaying an increased resistance to indentations, resulting in a higher measured HV.765Figure 5: The hardness values of several indents demonstrating the change in hardness based on region.Following confirmation of variance with fiber content, the fiber content of the horizontal line of indents was plotted as a function of distance from the left edge. As shown in Figure 6 there was minimal correlation between horizontal positioning on the cross-section. Furthermore, the increase of overall fiber content did not eliminate instances of low hardness values.

Furthermore, the increase of overall fiber content did not eliminate instances of low hardness values. This remained true when considering the vertical selection of indents.Figure 6: Vickers Hardness values of the horizontal indents for five samples of differing fiber content taken from the same printed transition.766Average Vickers HardnessWhile there was not an apparent trend in the measured HV in relation to location on the cross-sectional surface, there was an obvious trend for the average HV across all samples in each set. Beginning with the Non-mixing, ABS to CF/ABS sample set, Figure 7 shows the Average Vickers Hardness value and standard deviation of the indentations on each cross-section as a function of fiber content. As fiber content increases, the average hardness steadily rises from 13.2 HV to 16.4 HV, an increase of 24 %. Similarly, the standard deviation increases with increasing fiber content, rising by 300 % from 0.27 to 1.08.

Similarly, the standard deviation increases with increasing fiber content, rising by 300 % from 0.27 to 1.08. This indicates that, as expected, the increased presence of the fibers results in an increased resistance to deformation. Furthermore, the increased variation at higher fiber contents suggests that the poor mixing allows for regions of significantly different mechanical characteristics.Figure 7: The average Vickers hardness value and standard deviation of all viable indents for each sample.Figures 8a and 8b show similar trends for the other two sample sets: an ABS to CF/ABS transition printed using a screw geometry intended to improve mixing (Mixing ABS to CF/ABS) and a CF/ABS to ABS transition printed using the same non-mixing screw geometry as the first sample set (Non-Mixing CF/ABS to ABS). As seen in 6a, the average hardness and standard deviation still increase with fiber content (about 36 % and 270%, respectively) despite the change in screw geometry.

As seen in 6a, the average hardness and standard deviation still increase with fiber content (about 36 % and 270%, respectively) despite the change in screw geometry. While the maximum of 17.6 HV is a 7 % increase compared to the Non-mixing screw geometry, the maximum standard deviation also showed a small increase from 1.08 to 1.10. This indicates that the mixing geometry provided minimal improvements in mixing and producing a more homogenous internal morphology. Once again, the Non-Mixing CF/ABS to ABS transition shown in Figure 8b showed an increasing hardness value and standard deviation with increasing fiber content. A maximum 19.0 HV represented an 8 % and 15 % increase compared to the Mixing and Non-Mixing ABS to CF/ABS transitions, respectively. Similarly, the maximum standard deviation increased by 61 %.

A maximum 19.0 HV represented an 8 % and 15 % increase compared to the Mixing and Non-Mixing ABS to CF/ABS transitions, respectively. Similarly, the maximum standard deviation increased by 61 %. Within the CF/ABS sample set, average HV increased by 45 % whereas the standard deviation increased by 620 %.767Figure 8: A.) the average Vickers hardness for each cross-section from the Mixing ABS to CF/ABS sample set, and B.) the average Vickers hardness for each cross-section from the Non-Mixing CF/ABS to ABS sample set.Considering the increased maximums, the change in transition direction could be considered to increase both the variation and hardness of the printed transition. However, Figure 9 shows a comparison of all three sample sets as a function of fiber content.

However, Figure 9 shows a comparison of all three sample sets as a function of fiber content. This representation further indicates that an increased variation precedes any relevant increase in hardness when increasing fiber content, suggesting that there may be a critical level of fiber content in the sample before deformation resistance increases as a whole.Figure 9: The average Vickers hardness value and standard deviation for all studied cross-sections as a function of fiber content.768ConclusionsVickers Hardness of samples printed using the BAAM dual-hopper system is a useful characterization method for assessing the changes in mechanical performance surrounding the material transition region. The microhardness showed consistently different values based on the region of testing, i.e. carbon-fiber rich regions demonstrated a higher HV and greater resistance to deformation.

The microhardness showed consistently different values based on the region of testing, i.e. carbon-fiber rich regions demonstrated a higher HV and greater resistance to deformation. The average HV of a cross-section increased with increasing fiber content by 24 % to 45 % depending on sample set. Similarly, the standard variation of the average hardness increased by 270 % to 620 %. This behavior indicated that fiber-rich regions both increased deformation resistance and contributed to an increased variation throughout the specimen. Furthermore, this behavior was constant regardless of experimental conditions. The change from a Non-Mixing to Mixing screw geometry provided no reduction in variation, though there was a 7 % increase in maximum HV. There was also little difference created by switching transition directions.

There was also little difference created by switching transition directions. While the Non-Mixing CF/ABS to ABS group did display a improved HV and greater standard deviation by 15 % and 61 %, respectively, compared to the Non-Mixing ABS to CF/ABS, the overall trends remained the same.Future work will primarily focus on expanding the results of this study. A more comprehensive testing procedure will be used to investigate if increased sampling differentiates the Mixing and Non-Mixing geometries. The increased sampling will also be used to better analyze any positional trends in hardness within a single cross-section. Furthermore, a static mixing nozzle intended to improve blending of materials will also be utilized to print additional test specimens. Finally, a macroscopic characterization approach using Dynamic Mechanical analysis will also be included.AcknowledgementsResearch sponsored by the U.S.

Finally, a macroscopic characterization approach using Dynamic Mechanical analysis will also be included.AcknowledgementsResearch sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Industrial Technologies Program, under contract DE-AC05-00OR22725 with UT-Battelle, LLC.Thanks to Cincinnati Incorporated and Techmer PM for provided Material and Equipment. Further thanks to the University of Tennessee – Knoxville Joint Institute for Advanced Materials, the Metallography Lab group at the Joint Institute for Advanced Materials, the National Science Foundation for additional funding, and the University of Tennessee – Knoxville Center for Materials Processing for collaboration and additional funding.References[1] C.E. Duty, V. Kunc, B. Compton, B. Post, D. Erdman, R. Smith, R. Lind, P. Lloyd, L. Love, Structure and mechanical behavior of Big Area Additive Manufacturing (BAAM) materials, Rapid Prototyping J 23(1) (2017) 181-189. [2] T.

ARACTERIZING THE THERMAL-INDUCED DISTORTION OF LARGE-SCALE POLYMER COMPOSITE PRINTED STRUCTURESTyler Corum1, Johnna O’Connell1, James Brackett1, Ryan Spencer1, Ahmed Hassen2, Chad Duty1,21 University of Tennessee Knoxville, TN2 Oak Ridge National Laboratory Manufacturing Demonstration Facility Oak Ridge, TNAbstractThe Big Area Additive Manufacturing (BAAM) system has been used to print large-scale parts, such as automotive structures and molds for tooling, with fiber-reinforced polymer composites. Incorporating reinforcing fibers in printed parts is commonly used to increase stiffness and strength, but it also introduces significant anisotropy in the thermomechanical performance, which can lead to distortion and warping during thermal cycling. Characterizing and understanding how a printed tool distorts is crucial to maintaining tolerances and avoiding part failure.

Characterizing and understanding how a printed tool distorts is crucial to maintaining tolerances and avoiding part failure. This study uses digital image correlation (DIC) to measure the coefficient of thermal expansion (CTE) of a printed part from room temperature and to a known steady state temperature. The samples were printed with carbon fiber reinforced acrylonitrile butadiene styrene (CF-ABS). Various nozzle geometries were evaluated in this study with the intent of minimizing the thermal- induced distortion experienced by printed parts.IntroductionBackgroundAdditive Manufacturing (AM) has been a highly researched topic in academia and industry because of its ability to create complex geometries by adding material only where it is needed rather than removing unwanted material. Large-Scale Additive Manufacturing (LSAM) takes this a step further by offering the ability to print large (>1 m3) structures at high deposition rates (~50 kg/hr).

Large-Scale Additive Manufacturing (LSAM) takes this a step further by offering the ability to print large (>1 m3) structures at high deposition rates (~50 kg/hr). LSAM systems typically use fiber reinforced polymer (FRP) as feedstock material to decrease the coefficient of thermal expansion (CTE) and increase stiffness [1–3]. The lower CTE and increased stiffness also make FRPs a popular material for printed tooling applications. Not only is having a low CTE important but having a CTE mismatch between the tool and the created part may cause the final part to warp and fall outside of tolerances as shown in Figure 1.1921Figure 1. Mismatch of CTE values leading to part warpage is shown (adapted from [4]).While FRP feedstock can reduce CTE compared to the unreinforced “neat” resin, alignment of the fiber reinforcements during the printing process creates a highly anisotropic material. Thermomechanical anisotropy is introduced to the print by the shearing of fibers at the nozzle.

Thermomechanical anisotropy is introduced to the print by the shearing of fibers at the nozzle. The direction of these fibers influences the thermal properties [6–8]. This nozzle shearing causes fiber alignment along the print direction as shown in Figure 2. Fiber alignment caused by nozzle shear force aligns only the outermost fibers of the bead with the direction of the print but leaves the fibers located at the center of the print bead in a random, unaligned orientation [5–7,9]. The fiber alignment along the edge of the bead causes the printed structure to exhibit different CTE values at the edge of a bead compared to the center of the bead [6,7]. The difference in CTE values within each bead creates thermomechanical anisotropy across the entire LSAM structure.Figure 2. Fiber alignment due to nozzle shearing is shown above.Mechanical anisotropy is also introduced by the hierarchy of the AM process [1,2,10–12].

Fiber alignment due to nozzle shearing is shown above.Mechanical anisotropy is also introduced by the hierarchy of the AM process [1,2,10–12]. This is attributed to the layer-by-layer building technique utilized by AM shown in Figure 3. A single bead is extruded in the x-direction and is repeated, bead by bead, until the layer is complete. Layers are completed and build upon each other until the final structure is finished. Mechanical strength along a single bead (x-direction) is higher than across multiple beads (y-direction). Similarly, the strength within a single layer (x,y-direction) is higher than between multiple layers1922(z-direction) [1,10]. Other print parameters such as print orientation and raster angle can also effect the mechanical properties of the final part [13–15].Figure 3. Print hierarchy is shown above with the print beads (A) and print layers (B).Thermomechanical AnalysisCTE is traditionally measured using thermomechanical analysis (TMA).

Print hierarchy is shown above with the print beads (A) and print layers (B).Thermomechanical AnalysisCTE is traditionally measured using thermomechanical analysis (TMA). This method relies on heating a sample in a furnace, recording temperature with a thermocouple, and recording the material expansion using a linear variable differential transformer (LVDT) probe. TMA is a reliable and trusted method for measuring the CTE of a homogenous material, however, there are several challenges when applying this technique to LSAM structures. A TMA sample generally has dimensions on the order of 10 x 10 x 10 mm or smaller [16], but a single LSAM bead can have a width of 7 - 15 mm [1,2,7,17]. This size limitation of TMA means that the traditional method cannot accurately represent the large, LASM-printed structure using a single sample. The anisotropy of FRP structures complicate this further with the different thermomechanical properties across a single bead.

The anisotropy of FRP structures complicate this further with the different thermomechanical properties across a single bead. The limitations faced by TMA suggest that this method is not suitable for LSAM structures and other methods must be explored. Hoskins explored using Digital Image Correlation as an effective method for measuring LSAM samples without the size limitations of TMA [18,19].2D Digital Image CorrelationA common data collection method that is compatible with large parts is Digital Image Correlation (DIC). This technique is an accurate way to measure the strain of a part by imaging a surface before and after that part undergoes deformation [20–25]. A 2D DIC system uses a camera and computer program to track surface movement as shown in Figure 4. Surface movement is tracked by coating a sample surface in a white background and covering that background with small, black speckles. Once images are taken and compiled, a grid called a subset is imposed over each image.

Once images are taken and compiled, a grid called a subset is imposed over each image. The size of the subset is determined by the user and directly effects the accuracy of the results [22,24,26]. Selecting a subset too small may not provide the computer software with enough unique speckle points to track movement but selecting a subset too large may underestimate the strain experienced by the sample. With the appropriate subset over each image,1923the DIC software then compares each image to the first image taken. The first image taken in the set is referred to as the reference image. Comparing each image to the reference image allows the software to record how the position subset changed from the reference image to the current image. Within the subset, each pixel in the image is assigned a numerical value based on the grayscale contrast created by speckling and allows the software to compare position of the subset before and after deformation [20,23,25,26].

A position vector is then calculated from the center of the reference image subset to the center deformed image subset as shown in Figure 5. Each position vector in the image is used to calculate an average strain for that image. This process is repeated for each image until an average strain is determined for the set of images.Figure 4. Typical 2D DIC experimental setup is shown above (adapted from [23]).Figure 5. Subset tracking from the reference to deformed image is shown (adapted from [23]).DIC OvenThis research uses DIC to measure the thermal-induced distortion of LSAM structures printed with FRP using a unique DIC Oven setup. This method uses 2D DIC to track the strain experienced by a sample before and after thermal loading. The DIC recorded strain and known furnace temperatures can be used to calculate CTE. This method of measuring CTE is advantageous because it does not face the limitations of TMA.

The DIC recorded strain and known furnace temperatures can be used to calculate CTE. This method of measuring CTE is advantageous because it does not face the limitations of TMA. An experimental setup for the system used in this study is shown in Figure 6. Unlike in traditional 2D DIC with the camera horizontal to the sample, this novel setup has the camera placed under the sample. The location of the camera is important to consider when measuring the face of an expanding material. This camera placement minimizes expansion towards the camera from being read falsely as additional1924strain experienced by the sample. In this experimental setup, the speckled sample is placed on a fixed sample tray that only allows expansion away from the camera to reduce false strain readings. Initial studies of this novel system have been documented by Hoskins in previous work [18,19].

Initial studies of this novel system have been documented by Hoskins in previous work [18,19]. However, several improvements to the DIC Oven have been made for this study to reduce system noise (random, inaccurate values recorded by the system) and achieve more accurate results.Figure 6. A diagram (left) of the major components of the DIC Oven (right) is shown.System ImprovementsImproving the existing system was the first step in accurately measuring CTE of LSAM structures. The DIC Oven has been used in initial studies to measure the CTE of printed parts, but original results exhibited high standard deviations and had not been calibrated relative to a reference sample. This section details efforts to minimize the system noise by various DIC Oven improvements.Focus and ResolutionPositioning of the camera tripod was the first change to the original setup described by Hoskins [18,19].

The camera was repositioned so that the camera lens was 70 mm from the viewport and 250 mm from the sample. This reduced the amount of dead space in the images captured by the camera and helped the imaged sample maximize the field of view (FOV). With the new camera position established, the lens was refocused to improve image resolution, which allows the DIC computer software to more accurately track speckle location before and after distortion [21]. Figure 7 demonstrates the improved pixel resolution obtained by repositioning the camera. The 40 x 40 mm images shown in Figure 7 are the same sample, speckle pattern, and dimensions. The only difference in these images is the described changes in camera position and focus. The 2.5 x 2.5 mm images in Figure 7 show a zoomed in region with highlighted areas comparing the improvement in clarity from the left to the right image. Notice that in the before 2.5 x 2.5 mm image that this was the maximum zoom of the system.

Notice that in the before 2.5 x 2.5 mm image that this was the maximum zoom of the system. Maximum zoom after the1925improvements were made is also shown by the far right 0.81 x 0.81 mm image of Figure 7. This new maximum zoom demonstrates how clearly and precisely the camera can image each speckle, which was quantified by an increased in pixel resolution from 60 microns/pixel to 28 microns/pixel.Figure 7. Above shows the progression of the camera resolution before (left image set) and after (right image set) adjustments were made.Glare ReductionGlare from light entering the system and reflecting off floor tiles onto the bottom viewport during imaging as a reflection of the camera was also found to contribute to system noise. This glare reflection picked up by the camera is shown on the left side of Figure 8. To minimize this glare and get a more consistent image, the “Collar” was created and added to the camera lens.

This glare reflection picked up by the camera is shown on the left side of Figure 8. To minimize this glare and get a more consistent image, the “Collar” was created and added to the camera lens. The collar, shown in Figure 8, is a retrofit design made to fit the existing camera and block out any excess light or small particles that may interfere with the camera lens during imaging. The collar was modeled using SolidWorks and created using a fused filament fabrication (FFF) 3D printer. Once created and slightly adjusted to fit the existing camera lens, the collar was coated in a thin layer of Rust-Oleum High Temperature spray paint with a matte black finish to cover any existing sheen from the filament. Adding the collar significantly minimized glare seen by the camera, as shown by Figure 8, and provided the DIC computer software with a clearer, more consistent image.Figure 8.

Adding the collar significantly minimized glare seen by the camera, as shown by Figure 8, and provided the DIC computer software with a clearer, more consistent image.Figure 8. An image before (left) and after (right) the collar was added that shows significant glare reduction.1926LightingThe system lighting was also changed to evenly illuminate the sample. The original lighting within the furnace was a light bulb on opposite sides of the sample with a dimmer switch to adjust brightness. Lighting the entire furnace caused bright spots on either side of the sample from the lightbulbs themselves or from light reflecting off the stainless-steel walls inside the oven. Since photographers use indirect lighting to avoid taking an unevenly illuminated picture, the same logic was applied to the DIC Oven by adding angled light diffusers (ALDs) to the existing light sources. The ALDs are shown in Figure 9 and are constructed of aluminum foil, parchment paper, and staples.

The ALDs are shown in Figure 9 and are constructed of aluminum foil, parchment paper, and staples. This specific construction allows the ALDs to withstand the high temperatures in the DIC Oven without combusting or deforming. The diffusers help to light the inside of the furnace more evenly. Angling these diffusers allows the light to bounce off the bottom face of the furnace and indirectly light the sample as shown in Figure 9. The ALDs are also capped at both ends in aluminum foil to prevent light from bouncing off walls perpendicular to the light source. Figure 10 provides an example of how the change in lighting approach led to more consistent results from the DIC process.Figure 9. Shown above is the angled light diffusers (left and middle) and a lighting diagram of the furnace (right).Figure 10.

Shown above is the angled light diffusers (left and middle) and a lighting diagram of the furnace (right).Figure 10. An image taken before (right) and after (left) the ALDs were put into place.1927Miscellaneous ImprovementsThe furnace of the DIC Oven came standard with a viewport on the bottom and side. The side viewport of the furnace was covered by a sheet of aluminum to avoid any unwanted light in the system. The glass pane used in the viewport on the bottom of the DIC Oven was removed and replaced with quartz glass. Quartz glass is of better quality than standard glass and reduces the error caused by the inconsistent thickness and refraction of standard glass pane [20,22]. The shroud that surrounds the oven base, as shown by Figure 6, was redesigned so that it attached to each corner of the oven base by a small c-clamps, and the sides were lined with adhesive tape to create a permanent seal to prevent light from entering the system.

The design described by Hoskins included small dampeners located under each corner of the sample stage with intensions to further minimize vibrations [18,19]. However, this study found these dampeners had a CTE value of 345.3 μm/m°C and were removed to avoid any drastic expansion during DIC Oven imaging.Image Coverage and CountAdditional changes were made to the existing procedure to further reduce system noise. Speckle size and coverage is the key to tracking the position of subsets and will affect the accuracy of DIC measurements [22–24]. Speckles that are too large or too dense will not provide enough data points to the DIC software, while speckles too small or too sparse may not be unique enough to have the position recorded accurately. Room temperature tests were run with different speckle size and coverage since any strain at room temperature should equal zero. This meant results that minimized room temperature strain were most accurate.

This meant results that minimized room temperature strain were most accurate. It was determined that a 0.007” dot size that covered the sample with six passes by the Correlation Solutions Speckle Kit stamp provided contrast with the least system noise.The next step was validating the number of images necessary per image set to record all surface movement with the improved speckle coverage and pixel resolution. An image trial was performed at room temperature using 5 different samples to determine how many images were required in each image set to accurately represent the average strain experienced by a sample. Results of this trial are shown in Figure 11 where the x-direction continuous strain average (top plot) for each sample tested as well as the change in the continuous strain average (bottom plot) from the previous image. Only the x-direction values were presented because these values were ~10 times smaller than associated z-direction counterparts.

Only the x-direction values were presented because these values were ~10 times smaller than associated z-direction counterparts. The average strain of each image stabilized near image 22 for all 5 samples tested, suggesting that image 22 accurately captured the average strain experienced by each sample. This was further reinforced by the change in continuous strain plot that shows 97% of data points recorded after image 22 were within 0.1% change from the previous image. With 22 images accurately capturing the average strain of the image set, more than doubling the number of images taken to 45 will accurately capture the surface movement of the sample.1928Figure 11.

The Continuous Strain Average (top) and Change in Continuous Strain Average (bottom) show that 45 images are enough to accurately capture strain experienced by a sample.Validation of System ImprovementsData showed reduced system noise after the improvements were implemented, and theimpact of these changes are illustrated by the figures below. Images of the same speckled sample were captured at room temperature before and after any improvements were made. Vic-2D was used to create a 2D strain map of the before and after image set. These strain maps are shown in Figure 12. Since these images were taken at room temperature, the measured strain should be equal to zero. The before map shows that the system recorded a random spread of strain with a range of 365 microstrain to -1623 microstrain. Each image of the before image set was examined1929with no indication of a pattern present.

Each image of the before image set was examined1929with no indication of a pattern present. The after image of Figure 12 shows a more homogeneous spread of noise with an order of magnitude lower range of 182 microstrain to -186 microstrain. This remaining strain seen at room temperature by the system was considered general system noise that could be removed during post-processing of the acquired data. The minimized system noise was represented further by a graph that describes the microstrain recorded for each image at room temperature. Figure 13 shows the microstrain measured at room temperature for images captured before and after system upgrades. This chart shows the range of system noise was reduced by 96% and room temperature strain is now centered around zero as expected.Figure 12. The room temperature 2D strain maps above show reduced noise and range from before (left) to after (right) the system was upgraded.Figure 13.

The room temperature 2D strain maps above show reduced noise and range from before (left) to after (right) the system was upgraded.Figure 13. Shown above is a graph depicting room temperature strain recorded before and after improvements were put into place.1930CalibrationFollowing system improvements, the DIC Oven was calibrated to quantify accuracy and precision. The reference material chosen was 6061 Aluminum due to its homogeneity, defined CTE value, and the usage of aluminum as a calibration material for TMA systems [27]. Samples were taken from the center of a large aluminum cube (100 x 100 x 100 mm) to avoid any pre- stressed areas from the edges due to the material extrusion process. A TA Instruments Q400 TMA sample (10 x 10 x 8 mm) was taken to represent the x-direction with another sample of the same dimensions to represent y-direction.

A TA Instruments Q400 TMA sample (10 x 10 x 8 mm) was taken to represent the x-direction with another sample of the same dimensions to represent y-direction. A 20 mm slice of an xy face was also taken near the center of the cube for the DIC Oven, giving a 50 x 50 x 20 mm block that was spray painted white and speckled. The TMA values were treated as the target values for this calibration given the homogenous reference material and trusted TMA method for this application. Table 1 shows the resulting CTE values of the aluminum calibration from each measurement method. Each calibration sample was measured three times to track the accuracy and precision of the DIC Oven system against the TMA. The three runs for each sample were then averaged and the standard deviation was calculated for each direction of each method.Table 1.

The three runs for each sample were then averaged and the standard deviation was calculated for each direction of each method.Table 1. Calibration results are provided for the DIC Oven.Aluminum Calibration CTE Results (μm/m°C)Run 1 Run 2 Run 3 Avg Stn DevTMA Measurementx-direction 24.0 22.6 22.6 23.0 ±0.823y-direction 23.8 23.2 23.3 23.5 ±0.329DIC Oven Measurement y-direction x-direction 23.9 23.5 24.1 23.5 24.4 23.8 24.1 23.6 ±0.259 ±0.168Calibration results show that the DIC Oven produced similar results to TMA measurements. The average CTE of both systems is higher in the y-direction than in the x- direction. The average CTE recorded by the DIC Oven was within 1.0 μm/m°C of TMA for both the x and y-direction. The standard deviation of the DIC Oven was less than ±0.30 μm/m°C for both directions measured, which was lower than that measured by TMA.

The standard deviation of the DIC Oven was less than ±0.30 μm/m°C for both directions measured, which was lower than that measured by TMA. Results of the aluminum calibration suggest that the DIC Oven data represents the true CTE of samples.Experimental MethodsSample PreparationThe LSAM samples measured in this work were created using Big Area Additive Manufacturing (BAAM) using FRP feedstock [1,2]. There were 10 samples printed using BAAM made of 20% by weight acrylonitrile butadiene styrene (20% CF-ABS) that were prepped for measurement using the DIC Oven. There were 5 nozzle designs (standard and Design A-D) used to create these samples, and each design was used to print two samples for DIC Oven testing. The1931objective of each nozzle design was to slightly vary the orientation and distribution of fibers to create a different fiber alignment using each nozzle, but the specifics of each nozzle design are currently proprietary.

All samples were dried at 60°C for at least 6 hours. The samples were allowed to cool for at least 2 hours in a desiccator before speckling. Each sample was inspected so that the xy and xz faces with the least defects were selected for DIC. Having a parallel, defect-free face, is crucial for accurate DIC measurement [23]. The selected xy and xz faces were painted white using matte finish Rust-Oleum High Temperature spray paint. Best practice is to use as little spray paint as possible for the white background to avoid tracking the movement of the paint film rather than the sample surface itself [20]. Once the white background was dried, the black speckles were applied using the Correlated Solutions Speckle Kit. The 0.007” dot size was used in this study and each sample was coated in speckles 6 times for adequate coverage, as shown by the speckled sample in Figure 14. This speckle coverage was done for each xy and xz face of interest.

This speckle coverage was done for each xy and xz face of interest. TMA samples were also used in this study to compare methods. TMA samples were taken from the same printed parts and tested under the same parameters as DIC Oven samples. The TMA samples were 5 x 5 x 5 mm sections taken from the outer edge and the center of a single bead to provide a CTE value for each sample using both TMA and the DIC Oven.Figure 14. An image of a fully prepped sample is shown.Test ProcedureThe DIC Oven was used to measure the CTE values of each CF-ABS sample. Testing began by allowing the Correlated Solutions camera under the furnace to warm up for 10 minutes. A single sample was aligned on the sample tray as parallel as possible so that the x-direction of the physical sample was aligned parallel to the x-direction of the software. Lighting was adjusted so that the sample was lit enough for the camera to image the sample but not so bright that there1932was a glare on the sample.

Lighting was adjusted so that the sample was lit enough for the camera to image the sample but not so bright that there1932was a glare on the sample. Once the sample was properly aligned, brightness was set, and the camera was warmed up, the Vic-Snap software could be used to capture images of the sample. The room temperature was recorded, and a 45-image set was captured at room temperature at intervals of 100 ms. The furnace was then set to a temperature of 90°C and allowed to reach steady state for 6 hours. After the sample reached steady state, the camera was warmed up for another 10 minutes, and a second set of 45 images were taken at the steady state temperature of 90°C.Vic-2D was used to compile the 90 total images and determine the average strain value of each image using the DIC methods discussed earlier in the work. With the average strain known for each image, an average strain can be calculated for room temperature (𝜀𝑅𝑇) and steady state temperature (𝜀𝑆𝑆).

With the average strain known for each image, an average strain can be calculated for room temperature (𝜀𝑅𝑇) and steady state temperature (𝜀𝑆𝑆). Using the known conditions at room temperature (𝑇𝑅𝑇) and at steady state temperature (𝑇𝑠𝑠), the CTE of each sample could be calculated using equation 1. Any room temperature strain should equal zero and was treated as general system noise. Room temperature strain was subtracted from the steady state strain for this reason.𝐶𝑇𝐸 =𝜀𝑆𝑆−𝜀𝑅𝑇 𝑇𝑠𝑠−𝑇𝑅𝑇(1)Once all 10 samples were tested in the DIC Oven, the procedure was repeated with a “flipped” sample. Flipped samples were adjusted so that the x-direction of the physical sample was perpendicular to the x-direction of the software. This step was taken to minimize any possible biasing from lighting or from the sample alignment process. The flipped sample alignment was the only alteration to the procedure with all remaining steps repeated.

The flipped sample alignment was the only alteration to the procedure with all remaining steps repeated. The flipped orientation and normal orientation CTE values were averaged together based on the physical sample directions to determine a final CTE value of each sample for the x, y, and z-directions. The standard deviation of the CTE values was also calculated for each sample.The average CTE values measured by the DIC Oven were plotted for each direction. TMA data of the same sample was imposed over the DIC Oven values to directly compare methods. The TMA data was left as inner and outer bead values to compare CTE location to the average value measured using the DIC Oven. Since maintaining orthotropic properties is an important aspect for tooling applications and a driving force in this study the orthotropic ratios of the different nozzle designs were compared to that of the standard nozzle.

This was done by dividing the y-direction CTE by the x-direction CTE and was repeated for z divided by x. Normalizing each direction by x helped to better characterize orthotropic properties based on fiber alignment from different nozzle design.Results and DiscussionThe CTE values measured by the DIC Oven for each sample are discussed as well as the importance of these results with respect to tooling. There were 10 FRP samples testing using the DIC Oven. Sample sets were averaged to give a CTE experienced by the CF-ABS based on the1933nozzle used during printing. The data discussed in this section represents how CTE values of the samples were affected nozzle designs A-D compared to the standard nozzle to better determine how different fiber alignment effected the CTE.Standard NozzleThe CTE values in the x, y, and z-directions for the standard nozzle are shown in Figure 15. The trend of CTE data increasing from the x to the z-direction was expected due to fiber alignment.

The trend of CTE data increasing from the x to the z-direction was expected due to fiber alignment. The x-direction saw the lowest CTE because the fibers were most aligned in this direction and minimized the thermal expansion along fibers. The y-direction experienced high CTE due to the chaotic distribution of fibers allowing higher expansion across beads. The z- direction CTE was highest from both chaotically distributed fibers and weak bonds between layers. The standard deviation of the x, y, and z-direction CTE for the standard nozzle was low. TMA data from the same print was imposed over the DIC Oven data for a comparison between the two methods. CTE values calculated using TMA were taken from the inner and outer portion of the bead for each direction. Notice that the DIC Oven data falls within one standard deviation of the CTE values calculated using TMA.Figure 15.

Notice that the DIC Oven data falls within one standard deviation of the CTE values calculated using TMA.Figure 15. The CTE values of the standard nozzle are shown above with TMA valuesoverlaid for comparison.The orthotropic ratios for each nozzle are represented in this study by comparing CTE values between different directions. Figure 16 represents orthotropic properties of the standard1934nozzle by normalizing CTE measured in the y and z-direction to the CTE measured in the x- direction. Perfectly orthotropic properties would be represented in Figure 16 as a value of 1. The value of 3.41 for the y-direction represents that the CTE experienced in the y-direction is 3.41 times higher than the CTE measured in the x-direction. The standard nozzle produced samples with less orthotropic properties in the z-direction than in the y-direction.Figure 16.

The standard nozzle produced samples with less orthotropic properties in the z-direction than in the y-direction.Figure 16. CTE ratios normalized by the x-direction are shown for the standard nozzle.Nozzle Designs A, B, C, and DThe CTE relationship of each nozzle design between the DIC Oven and TMA are shownin Figure 17. There were no CTE values provided by Figure 17 to protect proprietary data, but the similarities between the DIC Oven and TMA on the CTE values can be shown and discussed relative to each other. There was no TMA data provided for Design C due to poor surface finish of the printed part. The CTE data collected follows the same expected trend of values increasing from x to z-direction as a result of fiber alignment due to nozzle shear. Notice that most DIC Oven measurements fall within one standard deviation of the TMA values. The only value not in this one standard deviation range, Design A’s z-direction, is close to having the standard deviations overlap.

The only value not in this one standard deviation range, Design A’s z-direction, is close to having the standard deviations overlap. Further, Figure 17 also shows several instances where DIC Oven results fall between the TMA inner and outer bead values. In other cases, the DIC Oven results were more similar to either the inner or outer value, reinforcing that TMA does not always accurately represent a LSAM sample just by averaging the inner and outer bead CTE values.1935Figure 17. The CTE values of the different nozzle designs are shown above with TMAvalues overlaid for comparison.Orthotropic properties of the different nozzle designs are shown in Figure 18. The orthotropic properties in the z-direction were improved by designs A-D when compared to the standard nozzle, however, all but one of the y-direction ratios were increased rather than decreased.

The orthotropic properties in the z-direction were improved by designs A-D when compared to the standard nozzle, however, all but one of the y-direction ratios were increased rather than decreased. These results indicate that orthotropic CTE properties can be influenced by nozzle design, and for this study, Design D provided the greatest improvement in orthotropic properties. Unlike the other three designs, Design D was able to improve orthotropic properties by lowering both the y and z- direction CTE ratios.Figure 18. CTE ratios normalized by the x-direction are shown for each nozzle.1936Relating Sample Properties to StructureVic-2D was used to create the strain maps shown in Figures 19 and 20. These 2D strain maps show a clear relationship between the properties of the structure and how the sample was printed. Figure 19 shows the xy-plane strain with homogeneous spread of low strain in the x- direction and a very defined purple strain in the y-direction.

Figure 19 shows the xy-plane strain with homogeneous spread of low strain in the x- direction and a very defined purple strain in the y-direction. There is a clear interaction between the beads on the top surface of the sample shown by the strain map. This was expected from fibers at the edge of the bead being aligned due to nozzle shearing. This purple strain line is a negative strain with high strain on either side of that line. This was interpreted as the bead interaction attempting to hold the structure together while the beads were clearly attempting to expand. The xy-plane results could differ when using a larger sample that captures more bead interactions. Figure 20 shows the xz-plane strain map. The xz-plane also saw a low strain value throughout the x-direction with very high, defined sections of red strain in the z-direction. The red lines in the z-direction represent high points of strain between print layers.

The red lines in the z-direction represent high points of strain between print layers. These were expected results from fiber alignment occurring at the nozzle to better orient fibers. This led to more stress experienced between layers. The 2D strain maps show clearly that there is a correlation between the sample properties and how the samples were printed.Figure 19. The 2D Strain map for the XY plane is shown with a color scale for reference.1937Figure 20. The 2D Strain map for the XZ plane is shown with a color scale for reference.Conclusions and Future WorkThe objective of this study was to improve the DIC Oven system and validate a method of measuring the CTE of LSAM structures. TMA is traditionally used to calculate CTE but has limited value when using large, anisotropic structures. The DIC Oven system was modified to accurately track surface movements of large AM samples as a function of temperature.

The DIC Oven system was modified to accurately track surface movements of large AM samples as a function of temperature. Significant improvements to the system reduced the noise level by 96 %, and the DIC Oven was calibrated to within 20 ppm accuracy using a known, homogenous reference material. The improved system was then used to characterize CF-ABS samples printed using BAAM with five different nozzle design. The CTE values from the DIC Oven were comparable to values collected using TMA taken from sections of the same print. As expected, the measured CTE values of the standard nozzle showed that the y and z-direction CTE were very different. More importantly, changing nozzle design to alter shear occurring in the nozzle successfully influenced the CTE of printed samples. Each design was able to decrease z-direction CTE, but only Design D was able to lower the CTE in both the y and z-direction.

Each design was able to decrease z-direction CTE, but only Design D was able to lower the CTE in both the y and z-direction. For large-scale printed samples, the DIC Oven provides a more accurate method of measuring CTE of LSAM structures than TMA as represented by the 2D strain maps that relate properties of the sample to the structure of the sample. Future work will focus on developing a method to predict the CTE of LSAM structures that enables part designs to reach correct dimensions after thermal-induced distortion occurs.AcknowledgementsThis research was funded by the Southeastern Advanced Machine Tools Network (SEAMTN) at the University of Tennessee, Knoxville and sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle, LLC.1938References[1] C.E. Duty, V. Kunc, B. Compton, B. Post, D. Erdman, R. Smith, R. Lind, P.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCHARACTERIZING THERMOMECHANICAL PROPERTIES OF LARGE-FORMAT PRINTED COMPOSITE POLYMER STRUCTURESTyler Corum1, Johnna O’Connell1, Maximilian Heres2, Jeff Foote2, Chad Duty1,31.) Mechanical, Aerospace, and Biomedical Engineering Department at the University ofTennessee, Knoxville, TN 379962.) Loci Robotics, Inc., Knoxville, TN 37902 3.) Manufacturing Science Division at Oak Ridge National Laboratory, Oak Ridge TN 37830AbstractLarge-format additive manufacturing (LFAM) is a manufacturing technique where a high volume of material is extruded in a layer-by-layer fashion to form structures that typically measure several meters in scale. The LOCI-One system is an LFAM-type system operated by Loci Robotics, Inc. that features a high throughput extruder mounted on a 6-axis robot arm.

The LOCI-One system is an LFAM-type system operated by Loci Robotics, Inc. that features a high throughput extruder mounted on a 6-axis robot arm. This research used the LOCI-One system to print single bead walls of 20% by weight carbon fiber reinforced acrylonitrile butadiene styrene (CF-ABS) at various layer deposition methods, print speed, layer times, and bead widths. The coefficient of thermal expansion (CTE) of the printed structures was measured to quantify effects of print conditions on thermomechanical performance. The CTE of the LFAM printed walls was measured using a large-scale digital image correlation system to characterize the distortion of the fiber reinforced composite material in the x- (print direction) and z- (between layers) directions.

This study determined that with varying print parameters the CTE measured in the x-direction was largely influenced by bead geometry with the CTE measured in the z-direction relatively unaffected by either the varying parameters or the method in which layer deposition occurred.IntroductionAdditive manufacturing (AM) is an area of interest for industry because this manufacturing technique gives the ability to create complex geometries by only placing material where it is needed. Large-format AM (LFAM) allows large amounts of material (>1 m3) to be deposited at fast rates (~50 kg/hr) [1–3]. Fiber reinforced polymer (FRP) is a common composite feedstock for these types of systems as it increases stiffness and reduces coefficient of thermal expansion (CTE) [1,4].

Fiber reinforced polymer (FRP) is a common composite feedstock for these types of systems as it increases stiffness and reduces coefficient of thermal expansion (CTE) [1,4]. During extrusion, however, the reinforcing fibers shear align in the direction of the print, which results in a highly anisotropic LFAM bead with a highly aligned shell and randomly oriented bead core or center [5–8] as shown by Figure 1. The reinforcing fiber resists expansion differently with lower CTE exhibited in the longitudinal direction (parallel to direction of deposition, x- direction) than the transverse direction (perpendicular to direction of deposition, z-direction). With fibers oriented based on location within the bead and these same fibers providing directionally dependent resistance to expansion, the final LFAM structure has complex microstructure with highly anisotropic thermomechanical properties [9,10].1747Figure 1.

Above shows the fiber alignment that occurs within an LFAM bead from nozzle shearing effects.Measuring CTE is traditionally done using methods such as thermomechanical analysis (TMA), but this fails to capture effects of the complex microstructure of an LFAM structure made using FRP [9–13]. TMA is a localized measurement technique that cannot accurately represent an LFAM structure both from sample dimension limitations and the small TMA sample must be representative of the entire part [9–11]. The method used in this study more accurately represents CTE by capturing global effects of the FRP microstructure by using 2D digital image correlation (DIC). This technique relies on speckling a face of interest and tracking position before and after thermal loading. Displacement vectors are then calculated and used to find the strain experienced by a sample face. This technique can capture large areas of interest and was used to better represent the LFAM structures in this work.

This technique can capture large areas of interest and was used to better represent the LFAM structures in this work. A DIC Oven setup was used in this study with a schematic of major components shown by Figure 2 below. This method uses speckling to track thermal induced distortion before and after thermal loading via 2D-DIC. Other work lists the calibration and validation of the DIC technique as a method to better represent LFAM structures [10,11].Figure 2. A schematic of the DIC Oven with major components is shown.1748Methods and MaterialsThe FRP material used in this study was 20% by weight carbon fiber reinforced acrylonitrile butadiene styrene (CF-ABS). Loci Robotics Inc.’s LOCI-One system is an LFAM type system that utilizes a single screw extruder attached to the end of a KUKA 6-axis robot arm to achieve high precision and speed during printing.

Loci Robotics Inc.’s LOCI-One system is an LFAM type system that utilizes a single screw extruder attached to the end of a KUKA 6-axis robot arm to achieve high precision and speed during printing. Single bead thick walls spanning only the x- and z-direction (XZ walls) were printed on the LOCI-One using different layer deposition methods and by varying parameters such as layer time, print speed, and bead width as shown by Table 1 below. The layer height was held constant at 5 mm for all samples. The layer deposition method used for each print was either a continuous layer deposition method (CLD) where the printer constantly printed each layer without interruption or the pause layer deposition method (PLD) where a dwell time was introduced to each layer to simulate a larger part being printed. To keep layer time consistent between “continuous” and “pause” printing, the print head velocity of the “pause” print was increased to accommodate a dwell period.

To keep layer time consistent between “continuous” and “pause” printing, the print head velocity of the “pause” print was increased to accommodate a dwell period. The flow rate of each sample was calculated using equation 1 below where flow rate was represented by Q and was calculated from the print velocity (v), bead width (w), and layer height (h). This value increased with higher print speeds as well as with a larger bead geometry. Once the XZ walls were printed, 2x2 inch samples were waterjet cut and machined. These small plaques were then dried and speckled using the technique in [11]. This sampling process of the XZ walls is shown by Figure 3.Q = vwh1749EQ. 1Table 1. Print parameters for this study are displayed below.

The printed XZ wall is shown by A, the machined DIC plaque is shown by B, and the speckled DIC Oven sample is shown by C.1750The CTE of the plaques was then measured using the DIC Oven technique to measure average strain across a known thermal load. This process began by warming up the camera, positioning the sample, and adjusting the lighting as described in [11]. The sample was then imaged at room temperature using a camera with the room temperature state represented by a 45 image set taken in 100 ms intervals. The furnace was then set to 90 °C and the sample was allowed to reach thermal equilibrium. Once steady state was achieved based on the thermocouple study performed in [10], the steady state images were taken in 100 ms intervals for a total of 45 images to represent steady state. The sample was then rotated 90 degrees and re-measured to alleviate biasing from the lighting or positioning as described in [11].

The sample was then rotated 90 degrees and re-measured to alleviate biasing from the lighting or positioning as described in [11]. The 90 total images were uploaded to Vic-2D where the average strain was calculated for each of the LFAM printed DIC plaques. With the known strain and temperature change, the CTE was calculated for each sample as in [10,11]. The average CTE of each sample was plotted ± 1 standard deviation.ResultsThe Vic-2D software was used to create a strain contour map of each sample as shown by Figure 4. The figure on the left shows a relatively homogeneous, low magnitude strain in the x- direction (strain measured in the horizontal direction) while the figure on the right shows discrete bands of high strain in the z-direction (strain measured in the vertical direction). This behavior followed the expected trend based on fiber alignment.

This behavior followed the expected trend based on fiber alignment. The CTE in the x-direction was much lower because this is the dominant direction of fiber alignment during extrusion, and the carbon fibers have a very low CTE along their axis. The z-direction shows red bands of high strain separating larger green bands of lower strain. The bands of high strain correspond to a higher concentration of fibers aligned in the x-direction that resist expansion poorly in the transverse (z) direction. The red areas correspond directly to the position of layer interfaces occur in the sample, where the most highly aligned fibers in the x-direction occur. The green, lower strain regions represent the more randomly oriented fiber at the center of each bead where a greater percentage of fibers are randomly aligned and thus provide greater resistance to thermal expansion in the z-direction.Figure 4.

The strain contour plot of sample 1S is shown here for the x and z-direction.1751The effects of aligned fibers are evident also in the Table 2 data which shows average CTE ± 1 standard deviation. Again, the x-direction CTE (CTEx) is an order of magnitude lower than in the z-direction CTE (CTEz) in each case regardless of print parameters. This CTE data is similar to data collected from thermomechanical properties of CF-ABS in [5,7].Table 2.

Below shows average CTE data ± 1 standard deviation with the corresponding flow rates for each sample.CTE Values Measured by DIC OvenLayer Deposition MethodSampleSpecimen CTEx (μm/m°C) CTEz (μm/m°C)Flow Rate (mm3/s)s u o u n i t n o Ce u s a P1234561S 1M 1L 2S 2M 2L 3M1* 3M2* 3L 4S 4M 4L 5S 5M 5L 6S 6M 6L10.0 ± 0.8 27.2 ± 1.3 29.3 ± 0.4 13.1 ± 1.9 28.7 ± 1.5 27.5 ± 0.6 22.5 ± 2.6 22.3 ± 1.6 27.8 ± 0.9 10.2 ± 1.5 23.7 ± 2.4 29.9 ± 2.3 12.4 ± 1.0 23.1 ± 1.5 28.6 ± 3.2 12.2 ± 2.1 24.1 ± 2.2 26.5 ± 1.2123.6 ± 6.6 126.0 ± 10.5 124.9 ± 4.6 124.8 ± 4.5 127.1 ± 5.2 124.9 ± 5.1 127.7 ± 8.1 117.9 ± 3.2 123.2 ± 4.7 123.1 ± 5.2 123.3 ± 4.4 123.5 ± 6.3 125.9 ± 5.9 117.1 ± 6.3 119.2 ± 5.1 123.5 ± 4.9 114.6 ± 7.7 118.3 ± 6.5746 1086 1416 375 607 722 250 248 367 3530 5162 6882 3578 5225 6750 3620 5203 6697The CTE of each sample with respect to flow rate in each direction is shown by Figure 5.

This was done to best compare effects of the varying print parameters on the CTE of each sample. Comparisons of the different layer deposition methods used in the study were made by comparing sample sets 1 to 4, 2 to 5, and 3 to 6. The plot in Figure 5A shows that CTEx decreases for the small bead geometry as the flow rate increases. This indicated that with increased material flow, more fibers were aligned in the x-direction and CTEx values decreased as a result. The large and medium bead geometries did not show this same trend because the larger beads had a bigger core with randomly oriented fibers. Also, the wider beads experienced more flow normal to the bead axis as the extruded material beneath the deposition head was “squished outward”. Figure 5C shows a trend of grouped data where the overall trend of the CTEx values increases from small to medium to large.

Figure 5C shows a trend of grouped data where the overall trend of the CTEx values increases from small to medium to large. The PLD data is grouped with respect to flow rate because print speed was held constant during PLD prints with the data plotted based on similar bead geometry. With respect to CTE axis, the groups indicate that CTEx values were not affected by the changing dwell times within each group of data points. For instance, with the small geometry of Figure 5C, the dwell1752times increased from 48 seconds to 108 seconds to 228 seconds with relatively no change in CTEx. The trend of increasing CTEx values again is driven by bead geometry and is attributed to ratio of the highly aligned bead shell to the randomly oriented centers. As the bead geometry became larger, the region of randomly oriented fiber at the center of a bead became larger while the highly aligned shell region remained relatively constant.

As the bead geometry became larger, the region of randomly oriented fiber at the center of a bead became larger while the highly aligned shell region remained relatively constant. The plot in Figure 5B plot shows that the CTEz values were relatively unaffected by the increasing flow rate. This makes sense as with the increased material flow, the layer height was constant, and did not change the amount of randomly oriented fibers relative to the highly aligned region. This trend is also seen by the Figure 5D plot that sees no relative change in the CTEz values as material flow rate increased regardless of the dwell time introduced with each bead geometry.Figure 5. The x-direction CTE is compared for both the CLD (A) and PLD (B) as well as z-direction CTE for both the CLD (C) and PLD (D).ConclusionsThe aim of this study was to better characterize the thermomechanical properties of LFAM structures made using varying print parameters by using the DIC Oven to measure CTE.

The strain plots presented in this work showed clear evidence of fiber alignment in the LFAM structures made of CF-ABS material printed using the LOCI-One system. Results showed that bead geometry was a dominant factor in the CTE values measured in the x-direction independent of the printing1753method (CLD vs PLD). As bead geometry increased, CTE in the x-direction decreased due to the larger amounts of randomly oriented fiber at the bead center relative to the highly aligned shell. The CTE values in the z-direction, however, were relatively unaffected by either the increasing flow rate or the introduced dwell times from the PLD. This effect was due to the constant layer height of each LFAM structure not changing the relative percentage of the cross section regions containing aligned or randomly oriented fiber.

This effect was due to the constant layer height of each LFAM structure not changing the relative percentage of the cross section regions containing aligned or randomly oriented fiber. Future work will be performed to relate the mechanical properties of these samples to varying print conditions as well as investigation thermal effects on layer bonding.AcknowledgementsSpecial thanks to the Southeastern Advanced Machine Tools Network (SEAMTN) at the University of Tennessee, Knoxville for funding this research. Machining of samples was performed by the MABE Maker lab at the University of Tennessee. This research was sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle, LLC.References[1] Area Additive Manufacturing (BAAM), 2015. https://doi.org/10.2172/1209207.

ARACTERIZING THREE DIMENSIONAL MICROSTRUCTURESFORMED BY PARTICLES: AN EXAMPLE OFELECTROPHOTOGRAPHIC PRINTING TONERDi Baia, Vineeth R. Patilb, Marcos Estermanc, and Shu Changa,b,*aCollege of Imaging Arts and Sciences bCollege of Science cKate Gleason College of Engineering *Corresponding Author Rochester Institute of Technology, Rochester, NY 14623-5603AbstractAdditive manufacturing of objects using powders can result in microscopic structures based on the properties of particles used. In additive manufacturing, piles of powder are consolidated by one or other mechanisms in order to form a part. The resulting microstructure within the part can impact its engineering performance. Predicting and controlling the engineering performance requires the characterization of the part’s material interior microstructures.

Predicting and controlling the engineering performance requires the characterization of the part’s material interior microstructures. H owever, obtaining the microstructure is difficult since the interior of a powder is a challenge to visualize.In this paper, unconsolidated powders are characterized. We have used Confocal Laser Scanning Microscopy to image three-dimensionally sediments of micron-sized poly-dispersed electro-photographic printing particles. Using image analysis tools, we have extracted each particle’s position and radius in selected sampling volumes. Through this methodology, we demonstrate the possibility to recreate the three-dimensional particle structures and extract values for morphological parameters for powder systems.

Through this methodology, we demonstrate the possibility to recreate the three-dimensional particle structures and extract values for morphological parameters for powder systems. The feasibility of direct and quantitative particle structural characterization can lead to much needed methods and tools to relate particle structures to the process of fabrication, the design of materials, and product performance in additive manufacturing.I n t r o d u c t i o nAdditive manufacturing technologies that use solid particles to construct parts share many similar issues with electrophotographic printing with solid toner. In both technologies, working with small individual particles gives a good flexibility in forming a part or an image from a bulk material. However, both techniques must comprehend how the individual particles will ultimately affect what is intended to be a continuous object.

However, both techniques must comprehend how the individual particles will ultimately affect what is intended to be a continuous object. The particles will converge into an ensemble with certain morphology after deposition.It is the particle morphology that determines the integrity of the ensemble and its material engineering performance. For example in the electrophotographic print (EP) process, the structure of the toner layer after deposition (the development stage) needs to be cohesive enough to successfully transfer from thephotoreceptor to paper under electrostatic fields and mechanical pressure without a significant change in structure to degrade the image quality. One would expect similar requirements for particle structures in the process of fabrication through additive manufacturing. The deposition of additional layers and the buildup of thermal energy can change the particle microstructure in a way that directly impact manufactured material quality.

The deposition of additional layers and the buildup of thermal energy can change the particle microstructure in a way that directly impact manufactured material quality. Creation of parts needs to meet certain engineering specifications, such as Young’s modulus, and requires controllable microscopic structures for the powder. T his microscopic structure depends not only on the particle material and size, but also on the particle delivery method.The characterization of particle structures has been primarily conducted by extracting microstructural parameters from simulations and comparing simulations to experimental inferences [1]. With recent advancements in imaging technology and image analysis tools [2-5], there has been an intensified interest in directly imaging the powder particle microstructures and extracting quantitative metrics [2, 6, 7].

A number of studies have demonstrated the feasibility of more direct and quantitative characterizations of particle structures based on 3D imaging [2, 3]. As previous experiments have a successfully identified the particle positions and size, the next step is to use this information to determine the morphological parameters of structures.It is the intention of this study to show the feasibility of quantifying powder morphology determined from 3D imaging. To demonstrate this feasibility, this initial study utilizes EP toner as the model material.

To demonstrate this feasibility, this initial study utilizes EP toner as the model material. EP materials have several advantages for this demonstration: (1) they fluoresce which enables the use of Confocal Laser Scanning Microscopy (CLSM), (2) the relationship between their particle morphology and print quality has been previously investigated [8, 9], and (3) there has been recent interests in the use of toner for additive manufacturing [10-15].With CLSM and image analysis tools, we have obtained coordinates and radii for particles in selected sampling volumes, recreated the three-dimensional particle structures, and extracted values for several morphological parameters for this powder system. The feasibility of direct and quantitative particle structural characterization can lead to much needed methods and tools to relate particle structures to the process of fabrication, the design of materials, and product performance in additive manufacturing.B a c k g r o u n d1.

Parameters for Particle StructuresThe microstructure in a powder depends on the material properties of the particles it assembled from, such as the inter-particle cohesion, the particle size and distribution, the particle material composition, and the particle surface properties. Powder particles can form a wide range of microstructures. These particle microstructures can range from amorphous and compacted hard-sphere systems (such as soil, grain, etc.) [4, 5] to highly open structures with dendritic tentacles (such as printing particles [8, 9], dust particles in outer space) [16].One parameter used frequently in describing a material’s microstructure is the packing fraction, which is defined as the ratio of the total volume occupied by all the particles in a cell to the volume of the cell. For cohesive particles such as EP toner, the structure’s packing fractions can be as low as 14% [8].

For cohesive particles such as EP toner, the structure’s packing fractions can be as low as 14% [8]. For large and non-cohesive particles such as grains, the structure’s packing fraction can be close to the random close pack limit of around 58% [17]. Sands with different sizes and distributions have been observed witha large porosity range. In a study by Al-Raoush et al., 31% to 48% were measured for four different samples [5].In addition to the packing fraction, other structural parameters include: (1) porosity, defined as the ratio of the empty space in the cell to the total volume of the cell, which is the supplement of the packing fraction; (2) local void ratio, defined as the ratio of the volume of the void space associated with the particle divided by the volume of the particle; and (3) coordination number, defined to be the average number of particles in contract with a given particle.

Much of the part’s structural parameters relate to the particles themselves through their materials properties, such as the inter-particle cohesion, particle sizes, and the particle size distribution.The determination of structural parameters needs to be carried out with a large enough sample volume so that the parameters become independent on the size of the sample [5, 18-20]. This minimum sampling volume, which is named “representative elementary volume”, [8, 9] signifies the scale of observation [5].An additional terminology that characterizes the void regions is the concept of an “excluded region” or an “available space” in the representative elemental volume [21]. The “available space” is the empty space that additional particle or particles can be inserted; while the “excluded region” is the space where the existing deposited particles prevent the addition of an additional particle.2.

Correlations between Materials Performance and Particle StructuresMaterial macroscopic performances depend on the structural parameters [22-34]. For example, the work of Gao, et al., indicates that the ductile fracture initiation toughness depends on the void volume fraction, the void shape, as well as the void distribution [23]. Gu et al. demonstrated that when energy density is insufficient, selectively laser-melted stainless steel will contain voids and untouched build particles in the material matrix produced [35]. The void region in amorphous (or random closed) particle packing has shown to be crucial to the behavior of particle systems [21]. Material properties have been observed as functions of the surface area and volume fraction of the available space [36].

Material properties have been observed as functions of the surface area and volume fraction of the available space [36]. For fluid flow through porous media, surfaces of the void region have been a subject of close investigation [37].There are also numerous studies indicating the dependence of the object’s performance on the properties of the particle from which it is composed. For instance, in the situation of amorphous packing, the packing density depends on the size distribution function [38]. In contrast, with cohesive particles, the packing density appears to be independent on particle sizes or the size distribution [8]. Frost [39] has reported that for two granular materials with the same void ratio, different preparation techniques could result in different mechanical responses. Chang et al. have shown that particles with similar compositions but different surface treatments could exhibit different packing and mechanical behaviors [9].3.

Chang et al. have shown that particles with similar compositions but different surface treatments could exhibit different packing and mechanical behaviors [9].3. 3D VisualizationThe ability to obtain particle positions in sample interiors through microscopic visualization is particularly appealing. The two main approaches have been Confocal Laser Scanning Microscopy (CLSM) and X-ray Computed Micro-Tomography (X-ray CMT).Confocal Laser Scanning Microscopy (CLSM) is a h igh-resolution optical microscopy with depth selectivity and is capable of imaging interior structures for samples that are transparent and fluorescent. Much work has been conducted in colloidal solutions to take advantage of similar refractive indicesbetween particles and the fluid for optimal imaging [2, 6]. The methodology involves the capture of stacks (in the depth direction) of planar images and then the analysis of the image stacks through their pixelated values.

The methodology involves the capture of stacks (in the depth direction) of planar images and then the analysis of the image stacks through their pixelated values. It has been shown that the particle positions can be tracked and reconstructed into a 3D volume space [2, 3, 45]. The particles used in colloidal systems are typically monodispersed [2].3D imaging with CLSM for micron-size particles in colloidal systems has contributed to substantial understanding on the atomic nucleation process and phase transitions [40-42]. Experimentations on these systems are performed with thin (a few layers of particles) and sometimes dilute samples for efficient light transmissions [43, 44]. The CLSM has also been employed for the studies of biological [42] systems. Included among the many CLSM applications for biological systems is the mapping of shape through the tracing of signals from fluorescence dyes that have been intentionally attached to the surface of the sampling objects.

There is a lack of explorations of using CLSM for granular powders.X-ray CMT has been used to study granular materials; primarily focusing on obtaining the representative distributions of local void ratios of porous media, evolution of the void ratio, localized deformations of granular materials [46, 47] and extraction of physically realistic pore network structure of porous media [48, 49]. Similar to CLSM, X-ray CMT captures stacks of images and performs analysis of images in pixelated values. D ifferently from CLSM, X-ray CMT does not require materials to be fluorescent. However, access to synchrotron radiation facility is needed in order to obtain high resolution imaging [50]. Much of studies of 3D particle mapping have been for much larger particles (~ mm scale) than are less interesting to additive manufacturing.Due to the nature of print-based manufacturing, the product design and its material selections are intrinsically intertwined and cannot be separated.

Well understood materials process-property-structure relationships can potentially lead to the establishment of a design framework that is unique to additive manufacturing. Although particle position tracking for colloidal and for millimeter-scale particle has been demonstrated, there is a lack of investigations for micron-size powder materials. This paper presents CLSM imaging of a model polydisperse micron-size powder material and the reduction of the structure it formed. In addition, this work reconstructs three-dimensionally the powder microstructures and extracts the powder system’s structural parameters with quantifiable metrics.3 D V i s u a l i z a t i o n a n d P a r t i c l e P o s i t i o n I d e n t i f i c a t i o n1. SamplesToner are typically 5-10 µm in diameters and consist of pigment (colorant), binders with both low and high melting point components, as well as charge- and flow-control agents (on the surface).

All three colors, cyan, magenta, and yellow appear to fluoresce.Yellow EP toner were used for this work and extracted from a HP3700 cartridge. The particle sediments were generated by dropping small quantities of toner onto a microscope glass slide using a spatula. Samples of sediments were enclosed with double-sided tapes and cover glasses on top to prevent toner contamination to the CLSM imaging system.The sample deposition was not a well controllable process and resulted in variations in sample heights.Analysis volumes were determined by selecting regions with similar sample thicknesses.2. Confocal Laser Scanning Microscopic ImagingThe CLSM imaging was conducted with a Leica TCS SP-5 Biological Confocal Microscope. The CLSM images were obtained with a 40X objective with Numerical Aperture of 1.1 and with water as the refractive medium to interface the objective lens and the sample cover glass slide.

The CLSM images were obtained with a 40X objective with Numerical Aperture of 1.1 and with water as the refractive medium to interface the objective lens and the sample cover glass slide. The sample was imaged in the fluorescence mode collected in the 475 nm – 610 nm range (with an excitation wavelength of 476 nm).The images were collected at a lateral increment of approximately 0.05 µm x 0.05 µm and an increment of 0.17 µm in the depth direction. Cross-sectional images (XY plane of 44.5 µm x 44.5 µm, roughly the spot size deliverable by a 600 dots per inch digital document printer) were obtained in stacks as a function of depths (Z-axis).The appearance of the “particle of interest” in the Z-position extends beyond the length of the radius measured in the XY plane. T he prolate spheroid manifestation is caused by the three-dimensional diffraction pattern of light emitted from a small “point” source (toner in this case) [2, 51].

T he prolate spheroid manifestation is caused by the three-dimensional diffraction pattern of light emitted from a small “point” source (toner in this case) [2, 51]. This effect is commonly observed and referred to as the actual image convoluted with a “point spread function” [2, 51]. The larger blurring in the Z direction gives the appearance of having more extent in that dimension. Although the effect causes elongation of the Z radii, the center position of the particle is not affected [51]. We assume these particles to be spherical based on observations of the particles in the XY plane. We have assigned the radius of each particle to be the maximum particle radius observed in the stack for that particle.Figure 1. An SEM image of toner particles.

We have assigned the radius of each particle to be the maximum particle radius observed in the stack for that particle.Figure 1. An SEM image of toner particles. Although most particles are approximately spheroids, there are particles taking on ellipsoidal shapes (A) as well as coalesced into agglomerates (B) (both encircled in the figure).The Scanning Electron Microscopy (SEM) images were taken for the purpose to validate particle sizes and distribution. SEM was performed with a JEOL-JSM-6100 system on toner particles covered with Au prepared through evaporation. Figure 1 displays a sample SEM image of toner used to characterize the size distribution. Figure 1 shows that most toner particles are spheroids. However, there are a few that are more ellipsoidal in shape (such as the one encircled and labeled A). T here also exist particles coalescing together to form agglomerates (such as the one encircled and labeled B).

T here also exist particles coalescing together to form agglomerates (such as the one encircled and labeled B). We excluded the ellipsoids and agglomerates from the size distribution estimation.3. Deduction of Particle Positions and Radii from CLSM ImagesOnce a stack of cross-sectional images have been collected, each particle’s (XYZ) coordinate and radius is identified through a four-step process: (1) select a particle to be the “particle of interest”, (2) measure the (XY) coordinate and radius for the particle in each of the cross-sectional planes, (3) identify the maximum radius and the corresponding Z-plane where the maximum radius has occurred, and (4) select the Z value corresponding to the maximum radius plane to be the particle’s Z-coordinate. The procedure was repeated for all particles in a selected sampling volume. Figures 2 and 3 illustrate this four- step process.Figure 2 depicts (a) the top view of a sample and (b) a magnified region with one cross-sectional view.

Figures 2 and 3 illustrate this four- step process.Figure 2 depicts (a) the top view of a sample and (b) a magnified region with one cross-sectional view. Fig. 2(a) shows that a 44.5 µm x 44.5 µm sampling specimens being boxed by a square in the upper left quandary of the image. This region is then selected for Z-stack imaging, as depicted with one view in fig. 2(b). Nine sampling specimens were examined.(a)(b)44.5 µm x 44.5 µm imaging areaFigure 2. (a) Image of an area of toner from which a small boxed area marked by white lines is selected for sampling in the depth direction. (b) One cross-sectional image from the sampling area where a particle of interest is analyzed for its XY coordinate (32.2 µm, 16.4 µm) and radius (4.3 µm) with the origin of axes being at the top-left corner.Fig. 2(b) displays a cross-sectional plane from the Z-stack of a “boxed” area.

2(b) displays a cross-sectional plane from the Z-stack of a “boxed” area. At this magnification, particles are visible and shown as grey areas (indicating fluorescence) through contrast against the black background (indicating no-fluorescence). The particles approximate spherical, with different diameters in appearance. Dark “cores” appearing inside of some of the particles shows the presence of a non- fluorescent material component within these particles. Conventional printing particles toners are known to incorporate more than one binder resins for the purpose of reducing the fixing/fusing temperature [52].Also displayed in fig. 2(b) are the measured X coordinate, Y coordinate, and the radius for the “particle of interest” for a “particle of interest”. This coordinate and radius determination is performed for all particles in fig. 2 (b) and for each of the cross-sectional image.Figure 3 illustrates process to locate the Z-position for the “particle of interest”.

2 (b) and for each of the cross-sectional image.Figure 3 illustrates process to locate the Z-position for the “particle of interest”. Displayed in fig. 3 are three cross-sectional images, (a), (b), and (c), corresponding to the same particle appearing in three different Z-planes. T his particle is marked in all three images with identical circles (in white). The complete stack referenced fig. 3 measures 44.5 μm by 44.5 μm by 27.2 μm in dimension. The Z positions for the three cross-sectional planes in fig. 3 are 4.0 μm, 9.2 μm, and 13.0 μm for figures 3(a), 3(b) and 3(c) respectively, as measured with reference to the glass cover slide on top of the sediment.As shown in fig. 3, with respect to the identical reference circle surrounding the particle, the particle appears in different sizes in three imaging planes as expected. Similarly, other particles also “appear” or “disappear” from each of the three views in fig.

Similarly, other particles also “appear” or “disappear” from each of the three views in fig. 3, as the cross sectional plane either intersects or does not intersect these particles.(b)(c)Figure 3. “Particle of interest” at three different image frames. The “particle of interest” is marked by the same sized circle in white. The figure shows: (a) appearance of fluorescence, b) emission at approximately full intensity, and c) diminishing of fluorescence. The largest radius of the particle occurs at around imaging frame (b). The imaging frame (b) provides the Z coordinate for the particle.By comparison, the radius for the particle in 3(b) is greater that the radii in figs. 3(a) and 3(c). The radii measured from different planes for the particle are compared to find a maximal value.

3(a) and 3(c). The radii measured from different planes for the particle are compared to find a maximal value. The Z position for the center of the particle is the Z-value for the maximal radius plane.The Z positions for the particle centers are also attainable through the maxima of the fluorescence intensity profiles. However, this method is not chosen here because of the dark “cores” aforementioned which obscure the true maximum in the intensity profile. In addition, the particle displayed in the three cross-sectional images appears prolate in the Z direction. This is from the optical effect discussed in the section of Confocal Laser Scanning Microscopy Imaging.

This is from the optical effect discussed in the section of Confocal Laser Scanning Microscopy Imaging. We have corrected this effect by using the radii obtained in the XY plane for the Z-direction for all particles.C h a r a c t e r i z a t i o n o f t h e M i c r o s t r u c t u r eBeing able to extract coordinates and radii for the particles means that the individual particleparameters and the film’s morphological parameters may be calculated quantitatively and directly.1. Size Distribution(a) CLSM MeasurementSize Distribution(b) SEM MeasurementSize DistributionFigure 4. Size distributions attained from diameters measured from (a) CLSM and (b) SEM images respectively. SEM radii are slightly larger than those determined from CLSM images.Figure 4 depicts the size distribution of toner particles with 4(a) measured from CLSM images and 4(b) measured from SEM images for the purpose of comparison and validation. With CLSM images, we evaluated 126 toner particles.

With CLSM images, we evaluated 126 toner particles. Figure 4(a) shows that the particles have a normal distribution with a mean radius of 3.4 µm and a standard deviation of 0.6 µm. The particle sizes range from 1.8 µm to 5.4 µm. Figure 4(b) shows a similar distribution analyzed for an order of magnitude more particles. The mean radius obtained from SEM images is 3.6 µm with 0.6 µm as standard deviation. The larger particle size range signifies the polydispersive nature of toner.Comparison between fig. 4(a) and 4(b) shows a slight difference in the mean radius as confirmed with a two-sample t-test. The smaller values from CLSM than those from SEM could be from diminishing fluorescence emissions at edges of the particles.2. Particle StructureFigure 5. Reconstructed structure of toner particles in a volumetric unit of 44.5 µm x 44.5 µm x 27.2 µm. The “suspended” particles are in actual contacts with and being supported by particles in the neighboring volume cells.

The “suspended” particles are in actual contacts with and being supported by particles in the neighboring volume cells. The structure is disorderly and open with much vacant space.Figure 6. The same reconstruction as fig. 5; however, with red spheres inserted to illustrate spaces available for insertion of other materials. I think this will rearrange once again when all the changes are accepted in the track changes featureUsing the XYZ coordinates and particle radii obtained, we are able to reconstruct the toner particle structures. An example of such a recreation is illustrated in fig. 5. The lighter portions of the spheres onthe outside of the reconstructed cell surfaces reflect the parts of particle volumes that do not contribute to the cell but are displayed here for visualization. The darker portions of the spheres represent the occupied space by the particles or the “excluded region”. Figure 5 shows that some particles appearing “suspended” in space.

The darker portions of the spheres represent the occupied space by the particles or the “excluded region”. Figure 5 shows that some particles appearing “suspended” in space. These particles are in actual contact with and supported by particles in the neighboring cells.Figure 6 exhibits the available space (in light grey) and the excluded particle volumes (in dark grey). For the purpose of illustration, the largest available volumes have been inserted with red particles to illustrate their possible dimensions. These vacant volumes represent regions that could potentially be used for material additions or become voids in the resulting material or product after post processing.3. Particle Packing FractionsIn Figure 5, t he particle cell reconstruction indicates very open structure and low particle packing fraction.

Particle Packing FractionsIn Figure 5, t he particle cell reconstruction indicates very open structure and low particle packing fraction. The total volume occupied by the particles within the volumetric unit can be summed up from particles within the sampling cell and partial particles on the cell surfaces, edges and corners.

Estimation from the particle measurement here has resulted in 13.3 ± 1.3 (%) packing fraction for the 9 samples examined. Comparable values between the result from this study and the predictions from simulations in the literature indicate that EP toners are cohesive with particle structures similar to that of the random ballistic deposition of cohesive particles.(1)(2)4. Coordination NumbersThe reconstruction in fig. 5 reveals the tentacular nature of the structure formed by toner. Figure 7 contains one connected chain lifted from the particle cell depicted in fig. 5. Figs. 7(a) is the sphere representation and 7(b) presents the vectors connecting contacting spheres with the spheres reduced to solid dots. The example in fig. 7 indicates that particles in this chain have a coordination number between 1 and 3.

The example in fig. 7 indicates that particles in this chain have a coordination number between 1 and 3. In this example particle cell, there are four terminating particles with coordination number of 1, five particles each contacting two other particles, and two particles each with three contacting particles. This tentacle results in an average coordination number of 1.8.(a)(b)Figure 7. A connected chain has been lifted out from the cell reconstructed in fig. 5. Displayed in 7(a) is the chain and 7(b) is its stick-diagram for vectors connecting the particles. 7(b) shows most particles have two nearest neighbors.The criterion to determine whether two particles are touching is the distance between the two particles being lesser than or equal to the sum of their radii. The coordination number illustrated with this particle chain is a characteristic of cohesive particles.

The coordination number illustrated with this particle chain is a characteristic of cohesive particles. In contrast, in the much studied random close packed sphere systems, the coordination number is approximately six per particle [53].C o n c l u s i o nWe have demonstrated the feasibility to determine particle positions in the interior of a powder, to visualize three-dimensionally the particle microstructure, and to quantitatively characterize the microstructure that the particles form. Using the Confocal Laser Scanning Microscopy and image analysis tools, we demonstrated this feasibility with EP toner in a sediment state. We have obtained and validated particle sizes and the particle size distribution observed through CLSM imaging. We have further employed this information to obtain the packing fraction of the toner system.

We have further employed this information to obtain the packing fraction of the toner system. In addition, we haveobtained examples of chain formations and the particle coordination numbers for a cohesive powder material.Although Confocal Laser Scanning Microscope is limited to only fluorescent particle powders, X-ray CMT can potentially visualize non-fluorescent or metallic powders, with resolution depending on the X- ray sources. In such cases, similar analysis can be conducted to obtain the particle parameters for metallic and non-fluorescent powders.The quantitative characterizations of the microscopic structures demonstrate direct measurement of particle structures. Such a measurement underscores a unique possibility to make significant progress in the understanding of the structure-property relationship for powder materials.A c k n o w l e d g e m e n t sThis research is supported by the Melbert B. Cary Jr.

Cary Jr. endowment by the College of Imaging Arts and Science and the Cross-Media Innovation Center at Rochester Institute of Technology as well as by the National Science Foundation Partnerships for Innovation Building Innovation Capacity (PFI: BIC) subprogram under Grant No. 1237761. A ny opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the views of the National Science Foundation. The RIT Confocal Laser Scanning Microscopy lab is supported through the National Science Foundation Major Research Instrumentation Program (#1126629), the RIT Office of the Vice President for Research, Kate Gleason College of Engineering, and the RIT College of Science. We would like to thank Ted Kinsman for capturing SEM images used in this study.R e f e r e n c e s [1] A. Hoffmann and H. Finkers, "A relation for the void fraction of randomly packed particle beds,"Powder Technology, vol. 82, pp.

218-221, Toronto, Ontario, Canada, 1998. [10] J. B. Jones, G. J. Gibbons, and D. I. Wimpenny, "Transfer methods toward additive manufacturing by electrophotography," in NIP 27: International Conference on Digital Printing Technologies and Digital Fabrication 2011, pp. 180-184, Minneapolis, Minnesota, 2011.[9][11] A. V. Kumar and H. Zhang, "Electrophotographic powder deposition for freeform fabrication," in10th Solid Freeform Fabrication Symposium, University of Texas, Austin, Texas, 1999.[12] D. Cormier, J. Taylor, and H. West, "An investigation of selective coloring with 3-D laser printing,"Journal of Manufacturing Processes, vol. 4, pp. 148-152, 2002.[13] A. V. Kumar and A. Dutta, "Electrophotographic layered manufacturing," Journal ofManufacturing Science and Engineering, vol. 126, pp. 571-576, 2004.[14] A. K.

4, pp. 148-152, 2002.[13] A. V. Kumar and A. Dutta, "Electrophotographic layered manufacturing," Journal ofManufacturing Science and Engineering, vol. 126, pp. 571-576, 2004.[14] A. K. Das, "An investigation on t he printing of metal and polymer powders using electrophotographic solid freeform fabrication," Doctoral dissertation, University of Florida, 2004. [15] J. Jones, D. Wimpenny, G. Gibbons, and C. Sutcliffe, "Additive manufacturing by electrophotography: challenges and successes," in NIP 26: International Conference on Digital Printing Technologies and Digital Fabrication 2010, pp. 549-553, Austin, Texas, 2010.[16] J. Blum and R. Schräpler, "Structure and mechanical properties of high-porosity macroscopic agglomerates formed by random ballistic deposition," Physical Review Letters, vol. 93, p. 115503, 2004.[17] J. M. Valverde, and A. Castellanos, "Random loose packing of cohesive granular materials,"Europhysics Letters, vol. 75, p. 985, 2006.[18] S. Edwards and D.

adding and Additive Layer Manufacturing with a laser supported arc processA. Barroi, J. Hermsdorf, R. Kling Laser Zentrum Hannover e.V., Hannover, 30419, GermanyAbstractThis paper describes the potential of a new process, combining the geometrical precision of a laser technique and the deposition rates of GMA cladding. Dilutions as low as 3 % can be achieved, leading to a high purity, in the first layer. Different material combinations like mild steel with X45CrSi9-3 are presented. Microsections for penetration depth determination show the high quality of the deposition layers. A hardness of the coatings of 63 HRC is reached. Hardfacing of shafts serve as an application example. The low heat input enables the process to build up structures. This results in a process variant for additive layer manufacturing which is also presented.

The low heat input enables the process to build up structures. This results in a process variant for additive layer manufacturing which is also presented. The production of macro-sized structures is shown and discussed.IntroductionMetal deposition techniques are used for cladding and form fabrication. Both applications deposit in layers, which are formed by single seams that are welded to the underlying surface. In cladding these seams are welded directly next to each other with a slight overlap. This forms an even surface. In Additive Layer Manufacturing (ALM) more complex geometries need to be handled and the process requires higher controllability. This controllability relates to switching the process on and off. A gap in the component can be produced when the process is switched off and a deposit is made when switched on. Using this method a component is built on a substrate.

A gap in the component can be produced when the process is switched off and a deposit is made when switched on. Using this method a component is built on a substrate. The substrate material can be the same as the deposition material, especially if the substrate is part of the component. If not it can be a cheap, easy to weld material which is removed afterwards. In cladding, where functional coatings are applied to workpieces, the materials differ. Another distinction of metal deposition techniques is the size of the components. In this work a new technique will be presented. It is suited for cladding and has the potential for fabrication, in both cases in the macro regime.Macro sized claddings and fabrications are presently done by various techniques. Plasma Transferred Arc for example is a process which provides extremely high deposition rates, but has rather high dilutions of 5-20 % [1]. Another technique is laser cladding, where high precision and almost no dilution are reached.

Another technique is laser cladding, where high precision and almost no dilution are reached. The major drawback is the need for high-power laser systems and the accompanying high costs. An approach to reduce these costs is Plasma Augmented Laser Cladding. Here, plasma is used to heat the deposition powder just below its melting temperature. A setback is the rather low deposition rate [2].Form fabrication experiments with gas metal arc welding (GMAW) and gas tungsten arc welding (GTAW) have been performed by Almeida et al. They were conducted with a titanium alloy as fabrication material. Applying the GTAW process, 1000 x 200 mm structures were built. The setback is the low deposition rate and a highheat input. Due to this heat input, the workpiece has to cool down to room temperature after each layer, resulting in a slow process. Cold Metal Transfer (CMT), a GMAW process with less heat input, has also been tested.

Cold Metal Transfer (CMT), a GMAW process with less heat input, has also been tested. The process uses a controlled current and a forward/backward oscillation of wire which is superimposing the wire feed. Despite this control the results with CMT show high penetration depth of around 2 mm [3]. To overcome this setback a laser enhanced cold GMAW process is used. The presented technique involves a GMA welding machine and a diode laser. The laser radiation is used for stabilization and guidance of the electric arc. Therefore, it is tuned to a wavelength of 811 nm by means of diode temperature control [4,5,6].ExperimentalThe setup is composed of a Merkle welding machine with ColdMIG technology. The torch is at an angle of 60° in the pushing direction. A 500 W direct-diode laser, with the beam inciding perpendicular to the workpiece (Fig. 1, 2), is used in order to stabilize the arc plasma.

The torch is at an angle of 60° in the pushing direction. A 500 W direct-diode laser, with the beam inciding perpendicular to the workpiece (Fig. 1, 2), is used in order to stabilize the arc plasma. The laser is pointing in front of the wire tip, with a slight overlap, to ensure that the laser radiation can interact with the plasma without melting the wire.Fig. 1: Sketch of setup alignmentFig. 2: Setup realizationCladding of LayersFor the cladding experiments, two different material combinations are used. Combination A consists of structural or mild steel as base material and 1.8401, a wear resistant steel, as deposition material. In combination B, 1.4718 hardfacing steel is used as deposition material and mild steel as substrate.Experiments with material combination A are performed on sheet metal. Therefore, an axis system is used to move the welding torch and the laser over the workpiece.

Therefore, an axis system is used to move the welding torch and the laser over the workpiece. After each seam, the welding head is moved back and aside by an offset sv in order to create a layer (Fig. 3). The seam offset is fitted to the track width br in order to achieve an even surface.Fig. 3: Geometrical layer parametersThe same principle is used for the hardfacing of shafts. A combination of a rotational axis and a linear axis is used to apply the deposition material as a single, helix type seam. The speed of the linear axis is matched to the rotation so that after each full turn the desired seam offset sv is reached.The important process parameters are the base current Ig, the lower limit current Ia and the average voltage U which the welding machine tries to maintain. Furthermore, the ratio of the wire feed rate WFR and the feed rate of the axis system FR determine the height and the width of the seam.

Furthermore, the ratio of the wire feed rate WFR and the feed rate of the axis system FR determine the height and the width of the seam. In all experiments a mixture of Argon and 18 % CO2 has been used as shielding gas.Additive Layer ManufacturingIn the ALM or form creation experiments, bar-like structures are produced using the 1.8401 steel. The mild steel substrate is not meant to be a part of the component in our tests. Its function is to carry the generated structure. The seam offsets in these experiments are perpendicular to the sheet plane and are determined by the thickness of the previous layer. They describe how much the welding head is lifted after a layer is welded. Therefore, in the first experiments, the height of the structure is measured after each layer.Experimental ResultsThe experiments show promising results in the cladding and the ALM process.

Therefore, in the first experiments, the height of the structure is measured after each layer.Experimental ResultsThe experiments show promising results in the cladding and the ALM process. With the same setup, it is possible to create protective coatings and to fabricate simple structures.Results of CladdingThe cladding experiments are performed in order to create a functional coating on a substrate with as little mixture of the materials and as low heat input as possible. For the material combination A, low currents together with a high welding speed show very little dilution.Table 1: Parameters for sheet coatingparameterIgIaUWFRFRsvvalue50 A20 A17 V3 m/min1.1 m/min1.3 mmFig. 4: Low dilution of base and deposition materialFig. 4 shows a cross section of a layer produced by these parameters (Table 2). Due to the low penetration depth of 0 mm to 0.17 mm, the dilution is very low and partly there is no welded bond.

4 shows a cross section of a layer produced by these parameters (Table 2). Due to the low penetration depth of 0 mm to 0.17 mm, the dilution is very low and partly there is no welded bond. The high purity of the coating leads to a hardness of around 370 HV (approx. 38 HRC).An extreme pushing angle of the GMA torch shows good results. Fig. 5 depicts a cross section of a layer obtained by this parameter set (Table 2). For a good welding quality at that torch angle, more power in the electric arc is needed. The progression of the penetration depth in the cross section shows a wavy behavior. This leads to pores on the substrate coating boundary at the interface of the welding seams. The coating does not show pores in the volume.

This leads to pores on the substrate coating boundary at the interface of the welding seams. The coating does not show pores in the volume. The reached dilution of about 10 % is higher compared to the previous shown results, but the surface has a better quality.Table 2: Parameters with extreme pushing angle of GMA torchparameterIgIaUWFRFRsvtorch anglevalue80 A40 A11 V3 m/min1.5 m/min1.25 mm30 °Fig. 5: Low dilution and even surface with material combination AExperiments with the parameter set in Table 3, using material combination B, result in a coating that is properly welded to the substrate, but has some pores between the seams at the substrate boundary (Fig. 6). The welding speed for the first seam was raised to lower the penetration depth. Despite this measure, it is still higher compared to the following seams. The dilution of the coating is as low as 3 %, with about 9 % dilution in the region of the first seam and less than 2 % in the rest of the coating.

The dilution of the coating is as low as 3 %, with about 9 % dilution in the region of the first seam and less than 2 % in the rest of the coating. This high purity of the layer leads to a hardness with an average of 63 HRC.Table 3: Hardfacing parameters for material combination B on sheet metalparameterIgIaUWFRsvFR, first seamFR, following seamsvalue85 A60 A19.5 V2.9 m/min2.5 mm0.6 m/min0.4 m/minFig. 6: Cross section of material combination BA transfer of the results from sheet metals to shafts is reached with nearly the same parameters (Table 4). A decrease of the voltage and the seam offset in addition to a slower feed rate of 0.4 m/min leads to an even surface (Fig. 7) and a properly welded deposit (Fig. 8). The hardness of the coating is high at an average of 63 HRC which provides excellent wear protection.Table 4: Parameters for shaft claddingparameterIgIaUWFRsvFRvalue85 A60 A17.2 V2.9 m/min2.25 mm0.4 m/minFig. 7: Seam helix with even surfaceFig.

7: Seam helix with even surfaceFig. 8: Cross section of shaft claddingResults for ALMThe ALM experiments are performed with the wear-protective wire material 1.8401. First results show that it is possible to build up structures by laying single seams on top of each other (Fig. 9).Fig. 9: Structure 90 x 27 x 6 mm, build by stacking single seamsIn these first tests it was already possible to reach an aspect ratio of the structure of 4:1. The results do not indicate a limitation for increasing this ratio. The heat flow from the welding area on top of the structure towards the substrate does not seem to decrease below a critical amount with the height of the fabricated component.The laser improves the process significantly. Fig. 10 shows a comparison of two structures. The structure on the left has been produced without laser stabilization and the structure on the right with laser stabilization. A significant misalignment of the seams results from this lack of stabilization.

A significant misalignment of the seams results from this lack of stabilization. An evenly erected layer-system is produced with the laser enhanced process. We assume that the laser helps to maintain the melting pool in the middle of the seam. Therefore, the melting is distributed evenly and does not flow exceedingly to one side.Fig. 10: Comparison of structures built with laserstabilization (right) and without (left)To build this structure we measured the height after welding each layer. This information is used to determine the correct height offset of the welding head for the next layer. To overcome the additional time needed for this measuring step we have analyzed the progression of the height of the structure with each layer. A linear fit to the resulting graph Fig. 11 shows very little deviation. Therefore, a fixed height offset was tested in later experiments.Fig. 11: Linear progression of the structural height with each layerAs can be seen in Fig.

11 shows very little deviation. Therefore, a fixed height offset was tested in later experiments.Fig. 11: Linear progression of the structural height with each layerAs can be seen in Fig. 12 problems arise at the beginning and the end of the structure when welding in only one, here pushing, direction. The middle of the structure shows a constant height, but since there is too much material at the beginning a build-up accumulates. The opposite is happening at the end of the welds. Not enough material is deposited and leads to a downward ramp that grows towards the middle of the structure with each layer.Fig. 12: Deposit build-up at beginning and lack of material at end of structureThe simplest way to correct for the volume issue at the beginning and the end of the structure is to weld bi-directionally. One problem is the non-symmetrical setup of the GMA torch. In previous experiments the change of the seam quality depending on the welding direction has been shown [7].

One problem is the non-symmetrical setup of the GMA torch. In previous experiments the change of the seam quality depending on the welding direction has been shown [7]. These tests have been repeated with at a slower welding speed. In Fig. 13, the seam quality, related to the welding direction, is compared. The dragging direction shows a distorted surface, whereas the surface in the pushing direction is smooth.Fig. 13: Comparison of welding directions dragging (top) and pushing (bottom)directionSince the bi-directional welding solution has such an appealing simplicity, we conducted further tests in two-layer systems. Two tests are performed, each without turning off the arc: a) welding in pushing direction, lifting the welding head and welding in dragging direction b) welding in dragging direction, lifting the welding head and welding pushing direction. The seams pictured in Fig.

The seams pictured in Fig. 14 both show smooth surfaces and have less volume issues at the beginning and the end of the structure.Fig. 14: Comparison of two-layer systems: a) first dragging, then pushing (top)b) first pushing, then dragging (bottom)The structure in Fig. 15 was fabricated with bi-directional welding. The bi- directional welding was done in one step with 3 min delay afterwards for cooling. No height measurement was needed. Due to the consistent layer thickness, an offset of 1.4 mm per layer could be applied. There is no build-up or lack of material at the ends of the structure.Fig. 15: Bi-directional welded structure with fixed height offset with a size of 260 x 26 x 6 mm (top)Ends of the structure without build-up or lack of material (bottom)Conclusion and OutlookWe have shown that the laser assisted process can be applied in cladding and also has the potential to be used in ALM. In cladding, it is possible to weld at high speed and low power.

In cladding, it is possible to weld at high speed and low power. This results in a low dilution, little heat input and a high hardness, due to the purity of the deposit.First tests have been performed in order to reduce the penetration depth and therefore the dilution to a minimum. The penetration depth is lowered that far, that there is almost no dilution with the parameters shown in Table 1. An average hardness of 370 HV (approx. 38 HRC) was reached in the coating. The setback with these parameters is the bonding. In the middle of each seam, the materials are welded but in between there is no strong bond.Tests with an extreme pushing angle (Fig. 5) lead to an even surface of the coating. The dilution is around 10 % due to the higher current that has been used. The few pores on the substrate coating boundary are small and should not have a negative influence on the quality of the layer, due to the properly welded bond.

The few pores on the substrate coating boundary are small and should not have a negative influence on the quality of the layer, due to the properly welded bond. A setback of the extreme pushing angle is that it limits the setup’s use in automated cladding of 3D contours due to a reduced free angle between workpiece and process head.The experiments with material combination B are performed with adjusted parameters (Table 3). There is a welded bond of the deposit and the substrate. Only in the seam-to-seam interface are there pores at the substrate boundary level. These will be far away from the surface after the finishing of the coating with grinding. The hardness of the deposit is measured with an average of 63 HRC providing a strong wear protection. A high purity of the layer is reached due to the dilution of only about 3 %. The first welded seam on the right side of the cross section (Fig. 6) shows a higher penetration depth.

A high purity of the layer is reached due to the dilution of only about 3 %. The first welded seam on the right side of the cross section (Fig. 6) shows a higher penetration depth. This is typical for starting seams and is already reduced by a 50 % higher welding speed as compared to the welds done previously. Therefore, the dilution in this region is higher at about 9 %, whereas the dilution of the following seams is less than 2 %. With slightly adjusted parameters (Table 4), it is possible to hardface shafts with helix type welding seam, reaching similar results as on the sheet metal.The achieved coatings are highly pure, so one layer will be sufficient where often, with conventional processes, two or more layers were needed. No pre-heating of the substrate is needed and the deposit reaches very high hardness for excellent wear protection.Future steps will be to investigate the 3D ability of the process. Therefore, experiments in different welding positions need to be performed.

Therefore, experiments in different welding positions need to be performed. Furthermore, a laser based adaption of the power distribution in the arc could lead to an evenly penetration depth, elimination of pores at the intersection of the seams and the substrate boundary.The ALM experiments show that the laser enhancement has advantages in the generation of structures. We assume that the laser helps to keep the melting pool in the middle of the structure. Therefore, the melting is distributed evenly and does not flowexceedingly to one side. Experiments with high-speed cameras will give information about this behavior. We have shown that structures with a high aspect ratio are possible. The 4:1 ratio achieved was only a first test. There is no indication that the ratio is limited by means of the process.

We have shown that structures with a high aspect ratio are possible. The 4:1 ratio achieved was only a first test. There is no indication that the ratio is limited by means of the process. The linearity of the structural height with the layers allows a good prediction of the needed height offset per layer, so no measurement needs to be done as an additional process step. Problems of uneven build-up at the ends of the structures have been overcome by bi-directional welding.The process has shown great potential for ALM but further research has to be performed. As a next step, layers with more seams in a cuboid form will be welded. The process ignition and stopping has to be optimized. Furthermore, seam height and thickness control by welding speed and wire feed rate have to be studied.

The process ignition and stopping has to be optimized. Furthermore, seam height and thickness control by welding speed and wire feed rate have to be studied. In order to control the heat distortion the heat management for the workpiece needs to be optimized.AcknowledgmentsParts of this research and development project is funded by the German Federal Ministry of Education and Research (BMBF) within the Funding Action “SME – Innovative: Research for Production (fund number 02PK2102) and managed by the Project Management Agency Karlsruhe (PTKA). The author is responsible for the contents of this publication.[1] Takano et al: Evaluation of processing parameters on PTA hardfacing surfaces,Welding International Vol. 24 pp. 241-247 2010[2] Wilden, J.; Emmel, A.; Bergmann, J.P.; Dolles, M.:Optimisation of energymanagement through Plasma-Augmented-LaserCladding (PALC); ITSC 2004: International Thermal Spray Conference 2004 pp.

viewed PaperCLAMPING CONCEPT FOR 6 SIDE HYBRID MANUFACTURINGS. Mischliwski1*, D. András2 and M. Weigold31 Department of Medical Technology Research, Center for Dentistry and Oral Medicine (Carolinum), University Hospital, Goethe University Frankfurt, Theodor-Stern-Kai 7, 60528 Frankfurt am Main, Germany 2 Saarstr. 22, 64342 Seeheim-Jugenheim, Germany 3 Institute for Production Management, Technology and Machine Tools (PTW), Technical University of Darmstadt, Otto-Berndt-Straße 2, 64287 Darmstadt, Germany *Corresponding: mischliwski@med.uni-frankfurt.deAbstractFor most technical applications, the surface quality and tolerances that result directly from additive processes are not suitable. Hybrid manufacturing as a combination of additive and subtractive manufacturing process steps can help solving this issue. In this work, a conceptional adjustable cast clamping process is introduced for a combination of Laser-based Powder-Bed-Fusion (LPBF) and milling.

In this work, a conceptional adjustable cast clamping process is introduced for a combination of Laser-based Powder-Bed-Fusion (LPBF) and milling. For component clamping during the milling process, the components are cast in place with a low-melting metal alloy, creating form-fit and force-fit connection. To prove the applicability, a rough estimation of occurring milling forces was conducted. In a subsequent series of tests, validation of clamping force was carried out using complex part geometries. A prototype fixture designed for this cast clamping process has been developed and tested. This fixture allows complex non-restricted 6-side machining of parts without moving it relative to the fixture or the need of any additional manual rework on part surfaces.Keywords: flexible clamping concept, 6-side machining, hybrid manufacturingIntroductionThe development of complex and sophisticated products demands continuously improving manufacturing processes.

Multiple factors can be relevant for the competitiveness and efficiency of a product during its lifecycle. The requirements for the product usually have conflicting goals. Customers demand for a tailored and technically mature product to satisfy the exact individual requirement, but for reasonable cost [1]. These conflicting goals are motivation for efficient manufacturing processes. In this work hybrid manufacturing is defined as a process combination of an additive and a subtractive process step. With a hybrid manufacturing process, it is possible to meet individual customer needs and create a valuable business case staring at a lot size of one. Hybrid manufacturing has been long recognized and is embodied in the availability of hybrid manufacturing systems with combined additive and subtractive capabilities to produce not only complex but also precise parts. It can increase the flexibility of the manufacturing process and reduce the amount of used material [2, 3, 4, 5].

It can increase the flexibility of the manufacturing process and reduce the amount of used material [2, 3, 4, 5]. The additive process step generates a semi-finished product, where the subtractive process step adds function and precision, e.g., geometrical accuracy and surface finish to the part where it is needed. The combination of a near net shape additive and subtractive manufacturing can reduce material and tool costs as well as overall machine time, resulting in cost advantages. However, combining two manufacturing process steps leads to the disadvantage of limited part design of both manufacturing methods. Machine processes are usually restricted to 5-side machining. A true 6-side machining is not possible without unclamping the part and clamp it in a new position after first machining steps.

Machine processes are usually restricted to 5-side machining. A true 6-side machining is not possible without unclamping the part and clamp it in a new position after first machining steps. After the re-clamping process, the parts usually must be repositioned and its position calibrated in order to minimize any possible error in rotation and translation with respect to the machine coordinate system. Especially for components with complex geometries, a lot size one and the size of dental products like crowns or bridges this re-clamping process is time consuming. Alternatives are custom-made clamping devices, e.g., negative clamping jaws, which, however, cause considerable additional costs for lot size one due to the individual manufacturing. These limitations inhibit a fully automated production process.

These limitations inhibit a fully automated production process. Common flexible clamping methods, like vacuum clamping (mostly for flat surfaces), clamping with formed clamping jaws1783or clamping with frozen fluids (thin layer of liquid on a flat freeze plate), already exists with the goal to solve this issue [6, 7, 8, 9]. However, none of these clamping methods are suitable for automated 6-side machining and handling of parts completely defined by free-form surfaces, like dental products. Depending on the geometrical shape, tools cannot reach every part surface and the lack of geometric reference during re-clamping challenges the position re-calibrating process. A promising approach shows a clamping by a low-melting metal alloy, though the most common used alloy contains Cadmium [10, 11, 12]. Acute inhalation of airborne Cd, e.g. from soldering or welding fumes, can cause severe chemical pneumonitis.

Acute inhalation of airborne Cd, e.g. from soldering or welding fumes, can cause severe chemical pneumonitis. Long-term exposure to low air concentrations can lead to chronic lung disease and possibly lung cancer. Long-term excessive exposure from air or food causes renal tubular dysfunction. In the case of oral intake, amounts as low as 300µg daily over a prolonged period are sufficient for chronic poisoning symptoms such as Itai-Itai disease [13, 14]. The acceptance in medicine and medical technology for toxic substances in the production environment is not present, even if it is only an auxiliary substance during production. The focus of this work is an adjusted clamping process chain using low melting alloy for mechanical 6-side postprocessing of via L-PBF additively manufactured parts.

The focus of this work is an adjusted clamping process chain using low melting alloy for mechanical 6-side postprocessing of via L-PBF additively manufactured parts. A proof of concept shows whether the clamping forces can be sufficiently absorbed during milling and whether components can be machined completely from 6 sides with sufficient accuracy.Clamping ConceptThe clamping concept developed in this work for the post-processing of components manufactured by means of L-PBF uses a Cadmium free liquified low melting metal alloy to clamp the part by cast clamping. The process was designed for the hybrid production of dental products such as crowns, bridges and abutments. In addition to a lot size of one and a small part size of about 15x12x12 mm³, dental products are characterized by a complete freeform surface, hence complicating an automated manufacturing process.

In addition to a lot size of one and a small part size of about 15x12x12 mm³, dental products are characterized by a complete freeform surface, hence complicating an automated manufacturing process. Therefore, a process chain and a clamping concept was needed which combines geometrical flexibility, the flexible handling and clamping of small and complex shaped parts. A process chain was designed (Figure 1) using low melting alloy as clamping mechanism that enables a mechanical postprocessing of all 6 sides without losing the relative component position and orientation inside the milling machine until the part is finished from all sides.

A clamping device (CD) that enables the handling of the part, the building platform, the casting process and the process of separation of part and building platform is implemented in the process chain.Figure 1: Process chain for hybrid manufacturing with low melting alloy as part of the CDFreeze clamping shows its flexibility due to the flat geometry of a freeze clamping plate when components show flat surface areas or many contact areas to the surface of the freeze clamping plate. In addition, due to the passive cooling of the component, thermal expansion or shrinkage must be considered and compensated if necessary. In casting clamping, the area of the component surrounded by the casting medium is made dependent on the casting height. This gives greater flexibility in positioning the component.

In casting clamping, the area of the component surrounded by the casting medium is made dependent on the casting height. This gives greater flexibility in positioning the component. The machining process takes place at room temperature.1784Development of a clamping device for 6-side machiningFor a fully capable 6-side clamping concept a clamping device is designed and the previously described process chain (Figure 1) is adjusted to necessary process steps [15]. The CD is shown in Figure 2. The base mount (1) is the center component of the CD. The part (2) is connected via support structures on a building platform (3). This building platform is pre-located and fixed by a rubber ring (4) and a mounting ring (5) to the base mount. A sensor (6) for measuring the poured liquid level height was integrated to the base mount. The system is completed with a cavity cover (7), which forms the bottom cavity for the alloy casting.

A sensor (6) for measuring the poured liquid level height was integrated to the base mount. The system is completed with a cavity cover (7), which forms the bottom cavity for the alloy casting. The base mount is directly screwed on a zero-point clamping system Polygrip Erowa ITS50 (not shown in Figure 2) that is used in the used 5-axis CNC mill Datron C5.Figure 2: Exploded drawing of the CDDue to the importance of the casting operation in the process, the CD was designed with geometries necessary for casting. In this context, the calculation values and methods commonly used in casting technology for widely used ferrous and non-ferrous metals could only be used to a limited extent for the design of an in-bi alloy casting process, since no material parameters are available here.

Casting diameter (12 mm), casting channel (1.128 cm²) and gate (0.7 cm²) were estimated according to table values for small components [16].Properties of clamping metal alloyThe following clamping concept uses a metal alloy with a melting point below 100°C. Therefore, the alloy can be liquified in hot water. A requirement for the alloy is a negative coefficient of thermal expansion (CTE) which results in an expansion after change of phase from liquid to solid state. Also clamping forces should not be lost over time, as it happens for example with base-, safe- or mellottes-alloy [17]. An Indium-Bismuth alloy (66 % / 34 % weight ratio) with a melting point of 72,4°C was chosen for the following experiments after a preliminary investigation [18]. Furthermore, this alloy exhibits the necessary negative CET after solidifying. Preliminary experiments showed low viscosity, good mold filling behavior and a low tendency to slag formation.

Furthermore, this alloy exhibits the necessary negative CET after solidifying. Preliminary experiments showed low viscosity, good mold filling behavior and a low tendency to slag formation. However, for the chosen alloy due to its lack of application, no data was found to describe the mechanical or thermal properties. The phase diagram is not fully confirmed and shows differences in the alpha and beta phase depending on the literature [19, 20]. Since only a eutectic alloy was used for the following1785experiments, these differences have no consequence. Material properties, like expansion/shrinkage over time and poisson’s ratio have to be determined.Experimental setupIn preliminary tests to measure the clamping force of the low-melting alloy, it was proven that the shear force load represents the critical case. The comparative force whose load the system has to withstand was chosen from a manufacturing process of dental products.

The comparative force whose load the system has to withstand was chosen from a manufacturing process of dental products. Conventional machining of dental crowns was chosen as the comparative process. In this process, machining with a 3 mm solid carbide ball endmill is identified as the biggest tool. The specific cutting force is determined according to Kienzle and Victor and amounts to 16.4 N [21]. Seven test specimens are arranged on the build platform. One is centered on the build platform. Six others are arranged at radius 16 mm and at 60° intervals around the central specimen. The distances were chosen so that there is sufficient safety distance to other specimens and the CD when using a 3mm ball endmill. The specimens are each cuboid with the dimensions 8x8x14 mm³ and have a 4 mm hole centrically along the Z axis. At a height of 7 mm, the test specimens have a circumferential notch which serves as a filling level for the casting process (Figure 3).

At a height of 7 mm, the test specimens have a circumferential notch which serves as a filling level for the casting process (Figure 3). The state after additive manufacturing and before mechanical postprocessing is the starting point. The specimen made of stainless steel AISI 316L (1.4404) is thoroughly cleaned and then measured on a Leitz PMM864 tactile coordinate measuring machine with a measurement deviation of 1.2+L/300µm (i.e., 1.4µm measurement deviation for a pitch circle diameter of approx.

60mm) at previously defined points with a ruby touch ball with the diameter of 1,5 mm (Figure 3).Figure 3: Left: measurement on a coordinate measuring machine; right: part on building platform inserted in CDThe measurement of the specimen was carried out over the processing steps along the process chain:not clamped, as build,  clamped, as build,  clamped, front side machined, component cast in, building platform separated.The calibration pins serve as reference points during the measurements on the front side until the casting process. The Pins are measured with 100 touches on the lateral side and 104 touches on the hemisphere. On each specimen, 100 touches of the drilling axis and 30 touches for each wall surface were performed to determine orientation and position. After the casting process and separation of the build platform, the same number of touches was performed from the rear side. L-PBF semi-finished products are subject to production-related tolerances.

After the casting process and separation of the build platform, the same number of touches was performed from the rear side. L-PBF semi-finished products are subject to production-related tolerances. In addition, the alignment of the specimens via the additively manufactured conical prepositioning pins can only be evaluated as a very rough component positioning. The specimens are measured indirectly within the milling machine via the calibration pins (Figure 4). The minimum machining oversize required must be determined experimentally to ensure the1786removal of all additive surfaces and thus geometric precision and surface quality of the milling process. Therefore, specimens with dimensions increasing in 50 µm steps, in the range 0 µm to 300µm, were manufactured.Figure 4: Left: calibration process in Datron C5; right: CD CAD visualisationThe machining process is divided into two different machining strategies:1.

Machining of all external surfaces of the specimen with oversize in 50µm increments from 0-300µm compared to the targeted geometry. Each side of the specimens is milled to half the specimen height. Where the bore has been milled directly to the internal diameter of 4 mm on all specimens to allow tactile measurement.2. Machining from the front beyond half the specimen height to a depth of 9 mm to get two machined surfaces on butt from both sides. The machining oversize has also been increased in 50µm steps in a range from 0 to 300 µm. The holes are milled equal to strategy one.After machining the front side, the CD is removed from the machine, the cavity cover is attached, and the specimens are cast in. After the alloy has cooled, the rubber and mounting ring are removed. The support structures, positions and calibration pins are manually sawed off with minimal force. The clamping device opened from the back is shown in Figure 5.

The support structures, positions and calibration pins are manually sawed off with minimal force. The clamping device opened from the back is shown in Figure 5. The CD is inserted back into the machine and the now revealed sixth side is now finish machined. The CD with the finished specimens is then removed from the machine. The specimens are heated to separate them from the alloy. The separated specimens are then cleaned from the alloy residues in an 80°C ultrasonic bath for 5 minutes. The height of the specimens is then measured using a dial gauge and compared with the height of the ideal geometry.

The height of the specimens is then measured using a dial gauge and compared with the height of the ideal geometry. Since there is no reference possibility in Z-axis after the casting and separation from the building platform, the height of the finished specimen plays a key role and can only be measured after the whole process is finished, including the separation from the alloy.1787Figure 5: Left: exposed specimens with partially left support structure; right: CD CAD visualization before (top) and after the separation of specimens and building platform (bottom) as of the starting point for machine finishing of the 6th sideResults and DiscussionThe experiment of the positioning accuracy has been carried out on one specimen within the scope of this work and is therefore only intended to provide a basic estimate of errors in the positioning accuracy of specimens along the process chain of a casting clamping system.

The measured values for the angular deviation and displacement are shown in Table 1. Specimens with an oversize bigger than 150 µm were milled completely without any as build surface from the additive manufacturing process left.Table 1: Averaged displacements and rotation along the process chainParametersAngular rotation Displacement x Displacement y Displacement zNot clamped, as build 0.0983° 0.0137 mm 0.0040 mm 0.0006 mmClamped, as build 0.0939° 0.0136 mm 0.0026 mm 0.0004 mmAfter removal of the building platform, 6th side not machined 1.305° 0.0422 mm 0.0129 mm 0.0012 mmThe measurement of the specimen heights after finish machining are shown in Table 2 with the used oversize. The results and deviation shown in Table 1 and Table 2 are sufficient to manufacture technical parts after general tolerance (± 0,1 mm) [22]. The deviation in Table 2 is a combination of the deviation of the CD, the zero-point clamping system and the machine repeatability.

The deviation in Table 2 is a combination of the deviation of the CD, the zero-point clamping system and the machine repeatability. Also, it is not clear what part of the deviation of the CD is caused by the alloy behavior or a possible lack of stiffness of the CD itself.

To determine the CD deviation, further experiments need to be executed.1788Table 2: Measured component heights and average deviationOversize0 µm50 µm100 µm150 µm200 µm250 µm300 µmTest series 114.249 mm 14.155 mm 14.063 mm 14.018 mm 13.794 mm 13.802 mm 13.655 mm214.231 mm 14.090 mm 14.065 mm 14.014 mm 13.798 mm 13.748 mm 13.634 mm314.263 mm 14.079 mm 14.074 mm 14.060 mm 13.852 mm 13.764 mm 13.638 mm414.231 mm 14.095 mm 14.055 mm 14.004 mm 13.922 mm 13.817 mm 13.656 mmaverage14.244 mm 14.105 mm 14.064 mm 14.024 mm 13.842 mm 13.783 mm 13.646 mmtarget14.3 mm14.2 mm14.1 mm14.0 mm13.9 mm13.8 mm13.7 mmaverage deviation 0.056 mm 0.095 mm 0.036 mm 0.024 mm 0.059 mm 0.017 mm 0.054 mmWhen measuring the specimens in the coordinate measuring machine, it has been shown that clamping the building platform (without casting) has no noticeable effect on the specimen position.

It should be noted, however, larger components have higher thermal stresses that can lead to distortion during the process, in particular during the separation of the components and the building platform. Measuring the specimens from the rear side has proved difficult because chips, support structures and the alloy melt can hinder tactile measurement (Figure 5). In addition, no reference geometry can be manufactured in the Z- direction, so that the component position can only be determined roughly from the rear side. The measurements have shown that the handling with the cast clamping system has the potential to be uncritical for accuracy in the process. However, an angular average deviation of approx.

The measurements have shown that the handling with the cast clamping system has the potential to be uncritical for accuracy in the process. However, an angular average deviation of approx. 1.3° of the specimens at the seven measured rectangular structures can be considered small (considering the component dimension of 8x8x14 mm³ and the small measuring depth of 5 mm), since the rough additive manufactured component surface and the small measuring depth can lead to larger deviations in the measurements. The casting process was carried out manually during the preliminary and main tests. A stable process, e.g., through automation, can increase the quality of the casting process. Nevertheless, all samples showed that the casting height of the solidified alloy even within a specimen sometimes fluctuates. It needs to be clarified whether an improvement can be achieved here through targeted heat management, e.g., in the form of fixed heating and cooling curves.

It needs to be clarified whether an improvement can be achieved here through targeted heat management, e.g., in the form of fixed heating and cooling curves. In addition, the specimens showed a consistent improvement in the surface finish at each 50µm oversize machining step. A completely bright machined surface was achieved with a selected allowance of 150µm. There was no noticeable offset at the abutting surfaces. No deviations were visually detected in the boreholes either. The finished specimens could be easily separated from the clamping alloy afterwards. Examinations under an SEM showed that all residues of the casting material could be removed through an 80 °C ultrasonic water bath. A critical point is the removal process of the building platform. The clamping forces of the alloy are limited, but not known due to unknown mechanical and thermal material properties.

A critical point is the removal process of the building platform. The clamping forces of the alloy are limited, but not known due to unknown mechanical and thermal material properties. Therefore, the widely used hacksaw machine for separating additively manufactured components and building platform cannot be used here, as the clamping forces would be exceeded. Nevertheless, there are approaches to keep the forces low by means of a suitable support structure or the usage of a torque-controlled saw.Conclusion and outlookWithin the scope of this work, a highly advanced clamping process and a clamping device tailored for hybrid manufacturing of complex part geometries and small lot sizes enabling automated 6-side machining is shown. The proof of concept was successfully made.

The proof of concept was successfully made. It is shown that clamping forces during milling can be sufficiently absorbed, and components can be completely machined from 6 sides with a maximum absolute deviation to the desired net-shape of less than 0,06 mm. However, the effect of the stiffness of the conical prepositioning pin structures used in this context may have to be investigated in further work.1789The optimization of the process chain and the clamping device has potential to minimize the occurring deviation and will be further investigated. Validating the limits of the process steps in the process chain as well as scaling would be a further step towards the mass adoption of additive manufacturing for lot size one technical components. The integration of all these different individual processes into one machine also poses a challenge under an economic point of view. Both the process chain and the CD are not limited to the L-PBF process used here.

Both the process chain and the CD are not limited to the L-PBF process used here. An application of other additive manufacturing processes with building platforms will be investigated. Also, a limitation to metal parts needs to be revisited. For other materials like high performance polymers (e.g. PEEK), the temperature of 80°C of this process must be uncritical.AcknowledgmentThis research received support from the Bundesministerium für Wirtschaft und Energie (BMWI) under grant ZF4016503References1. Kussmaul, R., Biedermann, M., Pappas, G., Jónasson, J., Winiger, P., Zogg, M., Ermanni, P. (2019). Individualized lightweight structures for biomedical applications using additive manufacturing and carbon fiber patched composites. Design Science, 5, E20. doi:10.1017/dsj.2019.192. Joshi, A., Anand, S.: Geometric Complexity Based Process Selection for Hybrid Manufacturing.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCLEATED PRINT SURFACE FOR FUSED DEPOSITION MODELINGC. S. Shafer, D. H. Siddel, A. L. Merriman, A. M. ElliottOak Ridge National Laboratory, Oak Ridge, TN 37831AbstractFused Deposition Modeling (FDM) has become popular among Additive Manufacturing technologies due to its low cost, speed, and geometric scalability; however, the primitive nature of the FDM build surface fundamentally limits the utility of FDM in terms of reliability, autonomy, and material selection. Currently, FDM relies on adhesive forces between the first layer of a print and the build surface; depending on the materials involved, this adhesive bond may or may not be reliable.

Currently, FDM relies on adhesive forces between the first layer of a print and the build surface; depending on the materials involved, this adhesive bond may or may not be reliable. Thermal contraction between the build plate and build materials can break that bond, which causes warpage and delamination of the part from the build surface and subsequent failure of the part. Furthermore, with each print, the user must use tools or special maneuvering to separate the printed part from the build surface as well as retexture or replace the used build surface. In this paper we present a novel build platform that allows for a mechanical bond between the print and build surface by using dovetail-shaped features. The first layer of the print flows into the features and becomes mechanically captivated by the build platform. Once the print is completed, the platform is rolled or flexed open to release the part from the mechanical bond.

Once the print is completed, the platform is rolled or flexed open to release the part from the mechanical bond. This design not only lowers the risk of delamination during printing but also eliminates the need for a user to reset or replace the build surface between print jobs. The effectiveness of each geometry was determined by measuring the distance at the pinch point compared to the distance that the extrusion filled below the pinch point. The Captivation Ratio was measured to compare the different geometries tested and determine which direction of extrusion creates a better ratio.IntroductionOver the course of development of Fused Deposition Modelling (FDM) Additive Manufacturing (AM), many improvements have been made to allow for a wide range of materials to be printed. The advancements that have been made to print surfaces, however, are each geared toward specific materials.

The advancements that have been made to print surfaces, however, are each geared toward specific materials. Considering the narrow range of print surfaces currently developed, it is unknown whether or not other materials may be compatible with the particular surface. This leads to one major limitation all FDM printers have, which is that materials can only be successfully printed if they are compatible with the build surface. Furthermore, build surfaces must be carefully monitored and maintained to ensure sufficient sticking during printing, which constitutes a significant operator burden. In order to adhere prints to the print bed, many of the build surfaces that are commercially available involve tapes, such as BuildTak™, Kapton® tape, and blue painter’s tape. Others involve glues, such as CubeStick™, large- and small-scale FDM printers still experience and WolfBite™. Unfortunately, delamination issues due to thermal changes and raster orientation (1, 2).

Others involve glues, such as CubeStick™, large- and small-scale FDM printers still experience and WolfBite™. Unfortunately, delamination issues due to thermal changes and raster orientation (1, 2). What is needed to solve this problem is a build platform that is compatible with all FDM materials, requires no replacing or resetting by the operator, and does not rely on adhesive bonding.This manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, worldwide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes.

The Department of Energy will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).1359The initial bond of the first print layer of an FDM print to the build surface is critical in achieving a successful print. In this work, we sought to design a build platform that is indefinitely reusable and will work with virtually any material by creating a mechanical bond between the print and a dovetail-shaped, cleated build surface. With this design the extrusion of the first printed layer flows into the grooves, mechanically bonding itself to the platform. There are currently no other systems on the market for this type of application; the majority instead rely on adhesive bonding. The United States Patent pending 14/548,701 includes this and any other technology that utilizes “recessive and protrusive” areas in order to captivate the printed parts (3).

The United States Patent pending 14/548,701 includes this and any other technology that utilizes “recessive and protrusive” areas in order to captivate the printed parts (3). The designs were created in order to create recessive areas that the extrusion would flow into and be captivated. The bond can only be released by bending the platform away from the part to open the grooves as shown in Figure 1.Figure 1: Removal of Captivated Polymer from Cleated Platform. This figure demonstrates how the build surface can be flexed to release from the printed part. Another design could be that each row is similar to links in a chain that can be rotated to release the printed part.ExperimentalTo test the effectiveness in captivating the first layer of an FDM print, cleated build surfaces of varying geometries were printed on a Fortus 900mc FDM printer out of ULTEM 9085; this material was chosen for its toughness and resistance to heat.

The cleated platform geometry is comprised of linear rows of “cleats” which are raised features that produce an overhang to capture the extruded plastic from the first layer of the FDM print. The cleat geometries tested were similar to a dove-tail groove geometry; however, one of the cleat edges is rounded. The rounded back of each cleat is representative of a future design where the cleats can rotate and release the printed part. The depth of the cavity between rows was approximately 0.5” deep, into which the extrusion would flow. These shapes were arranged in repeating rows in which all of the rounded edges were facing the same direction. Figure 2 shows a side view of two cleat profiles chosen for this study.

These shapes were arranged in repeating rows in which all of the rounded edges were facing the same direction. Figure 2 shows a side view of two cleat profiles chosen for this study. One cleat profile is a simple, one-sided dovetail groove with a varying angle (Figure 2a), and the other is a “wave” profile (Figure 2b).Figure 2: Cleat Surfaces printed via FDM (tan) in A) angled and B) wave profiles.1360The overall dimensions of the cleated surfaces were roughly 3” wide and 8” long, depending on the spacing between rows of cleats. Each printed cleat surface was affixed to the build plate of the Big Area Additive Manufacturing (BAAM) FDM printer, and Carbon Fiber-doped Acrylonitrile Butadiene Styrene (ABS) plastic was extruded over each surface.

The molten ABS was printed in both directions on the build surface in a “U” shape and flowed into the grooves in two different directions, as shown in Figure 3.Figure 3: Cleat geometries (tan) A) before and B) after polymer extrusion (black) with the BAAM system. The arrows indicate the direction of extrusion on the cleat surfaces.During extrusion with the BAAM, the molten plastic flows between the cleats, creates a foot-like profile, and solidifies upon cooling. These feet are what provide the mechanical bond to the print surface (See Figure 4). The critical features of the captivation are the width of the foot and the pinch point provided by the cleat profile. As the foot width exceeds the cleat pinch point, the extrusion becomes captivated.Figure 4: Side view of cleat profile with captivated polymer extrusion. Above is a side view of the build surface; the lightly colored section of the image is the cleated build surface and the dark section is the extrusion.

Above is a side view of the build surface; the lightly colored section of the image is the cleated build surface and the dark section is the extrusion. The pinch point can be seen between the rounded back of one row and the knife-edge of the adjacent cleat. The area below the pinch point is where the mechanical bond is formed. The “Foot Width” distance was used for ratio measurements.1361Multiple angles of the dovetail profile were tested for BAAM extrusion captivation,including 35°, 45°, and 55°. Two wave shapes were also tested with different spacings between rows. The cleat profiles, their dimensions, and the results of extrusion with the BAAM are given in Table 1.Table 1: Cleat geometries (tan) with polymer extruded (black) within vs. extrusion direction. This figure shows all of the geometries as printed on with the extrusion direction indicated.

extrusion direction. This figure shows all of the geometries as printed on with the extrusion direction indicated. In most cases, it is easily seen that there is a larger foot in the with direction than the against direction.GeometryExtrusion Direction: ← With CleatsExtrusion Direction: ←Against Cleats35° Angle, 0.225" Pinch Point45° Angle, 0.30" Pinch55° Angle, 0.14" Pinch55° Angle, 0.38" PinchWave Geometry, 0.28" PinchWave Geometry, 0.50" PinchAfter extruding over each cleated build surface with the BAAM system, the solidified extrusion (as shown in black in Figure 3 and Table 1) was removed by flexing the cleated substrate, and it was then measured with calipers. The size of the cleats and BAAM extrusion made measurement trouble-free. After measuring the “pinch-point” length of the cleated profile and the “foot widths” of each extrusion from each extrusion direction, the measurements were compiled.

After measuring the “pinch-point” length of the cleated profile and the “foot widths” of each extrusion from each extrusion direction, the measurements were compiled. The “captivation ratio” was calculated for each cleat and extrusion direction by dividing the foot width by the pinch point in each cleat. An example calculation would be as follows: If the pinch point at the top of the trough was measured to be 0.5 inches and the extrusion was measured to be 1.0 inches, then using an extrusion to pinch point ratio would result in 1.0”/0.5” which equals 2.0. An average of three feet of the extrusion was used in the ratio calculation for both directions of extrusion. The ratio data was compiled and displayed in a graph depicting the ratios of with and against the direction of the build surface for all of the variations, as shown in Figure 4.Results and DiscussionThe calculated captivation ratios were plotted against the pinch point widths as shown in Figure 5.

A trend can be seen in the plot between suggests that the captivation ratio decreases with an increase in pinch point width. Although the groove angle also decrease for two of these data1362points, comparing the two data points for the 55 degree groove angle at low and high values for pinch point further supports this trend. Further work is needed to isolate groove angle as a variable to discover its effect on captivation. Also an investigation into the relationship between the nozzle size and/or extrusion rate with the pinch point width is needed.Figure 5: Captivation Ratio vs. Extrusion Profiles Plot. This graph shows that there was a greater extrusion width in all of the build surfaces when extruded in the with direction, toward the rounded back first. The straight angled designs all fared better than the wave-shaped designs.

The straight angled designs all fared better than the wave-shaped designs. The wave shape has a greater directional dependence than the straight angle designs.Other trends from the plot include better captivation in extrusions that were deposited “with” the cleat profiles rather than against. This trend is significantly more pronounced for the wave profiles. When extruding in the direction of the cleat profile, the bead curls under the edge and is pushed farther down as the nozzle passes over, as indicated in Figure 5 by the white arrow. Conversely, in extruding against the cleat profile, the extrusion starts from the bottom of the cavity and is cut off by the overhang feature before it can push the plastic underneath it. Figure 6 illustrates the difference in flow behavior of the extruded polymer when extruded with and against the cleat profile.1363Figure 6: Flow of molten polymer between cleats during deposition.

Section (a,b,c) demonstrates the flow of plastic exiting the nozzle in the with direction, and section (d,e,f) demonstrates the against direction. It was hypothesized that the with direction would roll more plastic underneath the pinch point than the against direction due to the currents exiting the nozzle.Conclusion and Future WorkCleated build surface geometries were fabricated and tested for their ability to captivate amolten extrusion of thermoplastic as produced in the first layer of an FDM print. The cleat profiles were printed via FDM on a large scale and tested for captivation by the BAAM polymer FDM system. A captivation ratio was calculated based on measurements of the cooled extrusion and the cleat dimensions. The captivation ratio was plotted verses the cleat dimensions, and a general trend was seen between the captivation of the extrusion and the pinch point width of the cleats.

The captivation ratio was plotted verses the cleat dimensions, and a general trend was seen between the captivation of the extrusion and the pinch point width of the cleats. Further work is needed to explore the effect of cleat geometry (groove angle and wave or dovetail design) on the captivation of extruded polymer. Also, future work will consist of pull tests between printed parts and the build surface to more accurately characterize the captivation of extruded polymer in the cleats. Overall, a cleated build platform would provide a reusable and reliable build surface for FDM printers. This work is protected by United States Patent pending 14/548,701.AcknowledgementsThe authors would like to gratefully acknowledge Brian Post, ORNL for granting access to the Big Area Additive Manufacturing (BAAM) system that made these experiments possible.1364References1. T. Letcher, B. Rankouhi and S.

E CLEMSON INTELLIGENT DESIGN ENVIRONMENT FOR STEREOLITHOGRAPHY - CIDES 2.0Jill S. Kirschman Industrial Engineering Department jkirsch@ces.clemson.eduCharles F. Kirschman BPM Technology Inc.Georges M. Fadel Mechanical Engineering Department gfadel@ces.clemson.eduJoel S. Greenstein Industrial Engineering Department jsg@ces.clemson.eduLaboratory to Advance Industrial Prototyping Center for Advanced Manufacturing Clemson University Clemson, SC 29634 USAABSTRACTThere are a large number of commercial Rapid Prototyping (RP) devices available today. Allof these machines begin with a Computer-Aided Design (CAD) model, which is tessellated, sliced and then built layer-by-Iayer on the RP device. All of these operations, except the actual building of the part, are completed on a computer. Therefore, many improvements to the RP processes can be achieved through software, without affecting the RP devices or the warranties on them.

Therefore, many improvements to the RP processes can be achieved through software, without affecting the RP devices or the warranties on them. This has led to the development of a front-end software product to support the task of preparing the part to be built. The Clemson Intelligent Design Environment for Stereolithography (CIDES) is a user-centered interface between the CAD system and RP systems, primarily the Stereolithography Apparatus (SLA).CIDES 2.0 is designed to provide a variety of tools which are valuable to the users ofRP systems, including the ability to view and modify tessellated (STL) files, generate supports, and slice STL files into layer (SLI) files for use on an SLA. It also provides the ability to view SLI and merged (V) files. Furthermore, CIDES offers additional translation capabilities that make it valuable for other RP processes. The package has proven useful in the Laboratory to Advance Industrial Prototyping (LAIP) at Clemson University.

The package has proven useful in the Laboratory to Advance Industrial Prototyping (LAIP) at Clemson University. CIDES 2.0 is a new X Windows-based release based on the original version of CIDES with many additional features. A new Human Computer Interface is the major improvement to this release.INTRODUCTIONRapid Prototyping (RP) is the name for a number of processes that generate threedimensional objects directly from a Computer-Aided Design (CAD) solid model. Regardless of which process is used, the procedure to create these objects is to generate the CAD model, tessellate it, slice it, and finally build it on the RP device. All of these steps, except the actual building of the part, are completed on a computer. This was the motivation for the development569of a front-end software package to support the task of preparing a part to be built (Kirschman et aI., 1991A).

This was the motivation for the development569of a front-end software package to support the task of preparing a part to be built (Kirschman et aI., 1991A). The Clemson Intelligent Design Environment for Stereolithography (CIDES) is fundamentally a user-centered interface between a CAD system and the Stereolithography Apparatus (SLA).Rapid Prototyping is an active research area. Many researchers have custom-designed devices to conduct their experiments; however, much research is done using commercially available machines. Researchers can achieve significant changes to the parts produced on these machines without voiding the manufacturer's warranty by making changes to input files, or by interacting with the layer files that will be used to build a part. CIDES was developed for this purpose; it includes a number of features that have enabled Clemson University to enhance the capabilities of the RP systems that it has.

CIDES was developed for this purpose; it includes a number of features that have enabled Clemson University to enhance the capabilities of the RP systems that it has. CIDES provides input, output, and viewing capabilities for several common file formats. It also enables input file modification, support generation, slicing, and file format translation. A number of research activities have been based on the CIDES platform, including support structure generation (Kirschman et aI., 1991B), optimal orientation (Frank and Fadel, 1994), NURBS-based slicing (Vuyyuru et aI., 1994), adaptive slicing (Tata, 1995), and optimal placement of parts (Wodziak et aI., 1994). Currently CIDES is closely coupled with an SLA because that was the first available RP device, and the only one at Clemson University for several years.

Currently CIDES is closely coupled with an SLA because that was the first available RP device, and the only one at Clemson University for several years. However, the code is being expanded to work with other RP systems, including the selective laser sintering testbed at Clemson University (Terala, 1996).The original version of CIDES (version 1.0) was written in 1991. CIDES 1.0 runs on SunOS 4.3 or earlier. However, Clemson University has nearly completed the process of upgrading from this operating system to Solaris on all Sun workstations. Although the kernel of functionality is very portable, the CIDES 1.0 graphical user interface (OUI) and graphics engine were developed using the Hierarchical Objected Oriented Programming System (HOOPS) from Ithaca Software. HOOPS was an excellent choice for a graphics system in 1991, when there were few OUI standards for workstations. It is highly portable with versions for many platforms, and very powerful for display.

HOOPS was an excellent choice for a graphics system in 1991, when there were few OUI standards for workstations. It is highly portable with versions for many platforms, and very powerful for display. However, HOOPS has two shortcomings. First, HOOPS is designed for graphics and object display; it does not offer the OUI-building capabilities of a typical fourth generation language or a OUI builder. Second, Ithaca Software imposes royalties for the distribution of HOOPS-based software. As a result, CIDES 1.0 can not be freely distributed. This is the primary reason that CIDES has not found its way outside Clemson University, despite heavy use in the Laboratory to Advance Industrial Prototyping (LAIP).Another problem with the original version of CIDES is that it was designed for research purposes rather than general use. Very little attention was devoted to the user interface during the development of the product.

Very little attention was devoted to the user interface during the development of the product. Oould (1988) proposes that four principles of system of design be followed if one is to design a usable computer system. First, there should be an early and continual focus on the users. There should be direct contact with the users so as to understand their job and what they need to do their work better. Second, there should be integrated design. All aspects of usability should develop in parallel and under one focus or one person. Third, there should be early and continual user testing. To ensure that a design usable by and useful to its users, the system must be tested by its intended users. Finally, the design process must be570iterative. The system should be modified based on the results of user testing. Once the modifications are complete, user testing should be repeated.

Finally, the design process must be570iterative. The system should be modified based on the results of user testing. Once the modifications are complete, user testing should be repeated. This refine-and-test cycle is repeated until testing demonstrates that the product satisfies its usability specifications. The methodology used to redesign the CIDES interface applied these four principles.CIDES 2.0 was developed to achieve greater product accessibility. A new development platform, SOl workstation, was chosen and a new OUI was designed for the system. Several of the underlying data structures were revamped, and code was modularized to separate the functional kernel from the interface. This paper describes the redesign of the software from the top down and discusses the capabilities that are now available in the product.METHODOLOGYA four phase structured methodology was used to redesign the CIDES interface (Ulrich and Eppinger, 1995).

This methodology was chosen because it involves users in the design process, is user-centered and is self documenting. The first phase of the process begins with the development of a mission statement. For the CIDES redesign project, the mission statement was:Redesign the CIDES interface to make it more usable in and beyond the Laboratory to Advance Industrial Prototyping (LAIP) at Clemson University.Once the scope of the project was defined, representative users were identified based on several different criteria, including familiarity with CIDES, exposure to CIDES and similar packages, and frequency of use. A thirty minute semi-structured interview was conducted with five of the nine users. The remaining four users were interviewed via email because of their geographical location or scheduling difficulties. Upon completion of theses interviews, each user statement was translated into a user need statement.

Upon completion of theses interviews, each user statement was translated into a user need statement. To determine the relative importance of each need, the users were then asked to rate each need statement using a five point scale.The second phase of the design process translated the list of user needs into performance measures which were both testable and measurable. Some needs were not easily translated into quantifiable metrics. In such instances, metrics were created based on the user's subjective evaluations of the interactive system. Next, products similar to CIDES were identified and used as competitive benchmarks. The three products benchmarked were: the original version of CIDES, Partman (3D Systems) and Surfacer (Imageware). These products were benchmarked in terms of all the product performance metrics.

The three products benchmarked were: the original version of CIDES, Partman (3D Systems) and Surfacer (Imageware). These products were benchmarked in terms of all the product performance metrics. The information obtained from the interviews with users and the benchmarking process was used to establish a set of product specifications.Phase three consisted of concept generation and concept selection. It began with each designteam member individually generating ideas. Then the design team brainstormed together to create additional concepts and refine the initial ones. This process was repeated several times. The concept selection process was also iterative. First, the team selected the user needs and product specifications most able to differentiate among the concepts and prepared a selection matrix. The selection criteria were then placed in the rows of the matrix in descending order of importance.

The selection criteria were then placed in the rows of the matrix in descending order of importance. Next, the team chose one concept as a reference, and all other concepts were rated571on a +, 0, - scale against this reference on each criteria. When there was no difference, or if doubt existed as to whether a concept was better or worse than the reference, a score of "0" was given. The concepts were ranked based on the number of +'s, O's and -'s assigned to them. The decision to carry a concept forward was based upon team consensus, taking into consideration the rank, strengths and weaknesses of the concept. As a result, the team decided that a concept employing a Toolbar, Menu and Hot Keys would be carried on to the detail design and development phase.In the final phase, the selected concept was developed and tested.

Employing a process similar to that used by IBM to develop the 1984 Olympic Message System (Gould et aI., 1987), the team obtained feedback regarding the effectiveness of the interactive system being developed early and repeatedly. Initially the team used paper-and-pencil prototypes to test user reactions to the proposed design. Once the organization of tasks within the interface was completed the team began developing focused software prototypes. When the majority of the tasks were coded, the team moved to a comprehensive prototype so that users could identify any remaining design flaws before the overall design concept was frozen for detail design. Once detail design was complete, the interface was tested and refined using real tasks. Testing and refinement was an iterative process that began with the prototype developed in the detail design phase. Initial testing focused primarily on the layout of the interface as well as the labels used.

Initial testing focused primarily on the layout of the interface as well as the labels used. This was followed by four iterations of additional testing and refinement. During each of the four iterations, as many of the users as practical were included in the product testing. After the four iterations of testing and refinement, the interface appeared to be ready for final evaluation against the product specifications. The final version of the interface met or exceeded every product specification when it was tested in the Laboratory to Advance Industrial Prototyping (LAIP).STRUCTURECIDES 2.0 was redesigned to enhance portability and simplify improvements to the user interface. All of the core functionality of CIDES, including input/output, part manipulation, slicing, and supporting has been combined into a set of library routines that are called from the GUI (Figure 1). This separation enables the kernel, which is written in ANSI C, to be ported to different platforms more easily.

This separation enables the kernel, which is written in ANSI C, to be ported to different platforms more easily. Once the kernel is running, any interface from a simple text menu to a full graphical display or even a virtual reality helmet can be used on top of the kernel to provide access to the routines. The library contains routines to carry out all of the non-graphic functions capabilities discussed in the next section.Function CallsL-.,.-_...---.i+'---Data ----L---y-_~---lFigure 1. Kernel Structure of CIDES.This library approach enhances the utility of the software for research. While changes to the kernel must be more stringently tested, researchers can use the basic capabilities of input/output572and slicing without the risk of damaging the base code. Thus, additional user functionality can be added through software modules by anyone with a working knowledge of C.CAPABILITIESSince CIDES is a research testbed, its capabilities may change rapidly.

Its currentcapabilities can be broken into three categories: Display, Modification and manipulation, and Utilities. Each of these categories will be discussed in detail.The first function of CIDES is to display files of various formats. Included in this must bethe ability to read these files; currently STL, SLI and V files are supported. STL files are industry standard tessellated files used as input to all RP processes. CIDES supports reading and writing both ASCII and binary files that follow the STL file specification. Because some CAD packages generate non-standard STL files, CIDES is capable of handling several common anomalies, including non-standard headers or lines, incorrect case or number format, and bad concatenation. STL files can be viewed from any of the standard views (Figure 2) or the user can pan, rotate, and zoom via mouse commands.Figure 2. STL Viewing Screen.Currently CIDES supports only flat shading of the image.

STL Viewing Screen.Currently CIDES supports only flat shading of the image. Flat shading helps the user to detect errors in the STL file, especially when combined with what is called "disco lighting" by the users. This type of lighting is simulated by placing four different color light sources at the comers of a cube surrounding the part to give the part a multicolored appearance. Since the color573of a facet is dependent on its normal, any errors in the normal of a facet are quickly evidenced by a triangle of one color in the midst of a face that is a different color.The second type of file read and written by CIDES is the SLI file, which is a 3D Systems Inc. slice or layer file. The dialog which allows the user to interact with these files is shown in Figure 3. SLI files are displayed layer-by-Iayer, allowing the user to turn on and off various vector types.

The dialog which allows the user to interact with these files is shown in Figure 3. SLI files are displayed layer-by-Iayer, allowing the user to turn on and off various vector types. For more rapid viewing, the user can choose to run up or down through the files, selecting the delay before each layer is replaced on the screen via the slider; delays range from no delay to half of a second. As with the STL file, the user can pan, zoom, and rotate with the mouse. The third type of viewable file is the V file, which are 3D Systems Inc. "merged" file used to build a part. These are also layer files, and are viewed using the same tools as the SLI files.Figure 3. Layer Viewing.CIDES also has several capabilities that help the user modify and manipulate files. Objects can be moved, rotated, and scaled in order to achieve the best build.

Layer Viewing.CIDES also has several capabilities that help the user modify and manipulate files. Objects can be moved, rotated, and scaled in order to achieve the best build. The move feature includes both Move, which translates the part a relative distance along any of the coordinate directions, and a Move To, which sets the part minimum to a specified location. The latter is useful in order to assure that the part is in the positive octant. Another modification function sorts the facets of the STL file based on height for faster slicing.The utility functions provide the user with capabilities to improve the SLA output or input new files. Verification routines check for and fix inverted normals and provide information on574missing facets. The support algorithm (Kirschman et aI., 1991B) allows the user to customize a support structure and view it on the screen.

The support algorithm (Kirschman et aI., 1991B) allows the user to customize a support structure and view it on the screen. The slicing routines are not as robust as their commercial variants, but the support can be sliced when generated, and a variety of hatch and fill capabilities are available. Volume information is provided for cost estimation. Algorithms are also included to translate from AutoCad neutral (DXF) files to STL files. CIDES offers sufficient capability to help the user of an RP system get more out of the system. The results of current research at Clemson University will be incorporated into future releases of CIDES to increase the utility of the software.CONCLUSIONSRapid prototyping technologies are used to produce parts quickly directly from their CADrepresentations. However, many areas of the technology have not been made accessible to researchers wishing to "extend the envelope" ofRP technology.

However, many areas of the technology have not been made accessible to researchers wishing to "extend the envelope" ofRP technology. CIDES includes a number of important features that have enabled Clemson University to increase the capabilities of the Rapid Prototyping (RP) systems that it has through software. CIDES provides input, output, and viewing capabilities for several common formats. It also provides facilities for STL file modification, support generation, slicing, and file translation. This has made CIDES a research testbed for a number of different projects at Clemson University. Currently CIDES is closely coupled with the SLA because that was the first available device, and the only one at Clemson University for several years. However, the code is being expanded to work with other RP systems, including the selective laser sintering testbed at Clemson University.

However, the code is being expanded to work with other RP systems, including the selective laser sintering testbed at Clemson University. CIDES 2.0 incorporates a new GUI which was created using sound GUI design principles.ACKNOWLEDGMENTSThe authors thank NASA for its support of this research through NASA grant NAG3-2046. The authors also wish to acknowledge the contributions of Oliver Heim, Shannon Hughes, and Vishal Jadhav to the development of the CIDES 2.0 interface. Finally, the authors would like to thank all the users who were involved in the development and use of CIDES; their insights, comments and suggestions were the driving force of this project.REFERENCESFrank, D., and Fadel, G. (1994). "Preferred Direction of Build for Rapid Prototyping Processes."Proceedings ofthe Fifth International Conference on Rapid Prototyping Dayton, OH.Gould, J. D. (1988). How to Design Usable Systems. In M. Helander (Ed.), Handbook of Human-Computer Interaction.

inical Evaluation of Prosthetic Sockets Manufactured By Selective Laser SinteringBill Rogersa, Andrew Gittera, Gordon Boskera, Mario Faustinib, Mahendra Lokhandeb, Richard CrawfordbaThe University of Texas Health Science Center at San Antonio bThe University of Texas at AustinAbstractA pilot study was undertaken to evaluate the clinical acceptance of prosthetic limb sockets manufactured using solid freeform fabrication (SFF). The fabrication of sockets for amputees is a natural application for SFF. The socket is the part of the prosthetic limb that fits onto the amputee’s residual limb. Each socket is custom manufactured for each individual amputee. Four amputees were successfully fit with sockets created using selective laser sintering. The scope of the study included software development, finite element analysis, materials testing, and clinical evaluation.

The scope of the study included software development, finite element analysis, materials testing, and clinical evaluation. This paper discusses socket design issues and clinical testing results.PurposeThe purpose of this pilot study is to demonstrate the feasibility of Solid Freeform Fabrication (SFF) using Selective Laser Sintering (SLS) to fabricate variable compliant wall prosthetic sockets for trans-tibial amputees. The study includes an engineering design component and a clinical case controlled comparison of variable compliant wall prosthetic sockets with conventionally fabricated prosthetic sockets.The overall long-term objective of this project is to develop a system to fabricate variable wall compliant prosthetic sockets directly from digital residual limb shape information using freeform fabrication techniques.

These sockets are expected to improve prosthetic socket comfort, improve tissue-loading characteristics, reduce skin discomfort and breakdown, and reduce fabrication time socket and cost fabrication techniques.compared to conventionalIntroductionlimb amputees in the United States with 60,000 new major lower extremity amputations performed yearly. The majority of these amputees are fit with a prosthetic limb as part of their rehabilitation and return to independence. Because of changes in the shape and volume of the residual limb following amputation, the typical amputee will require a new prosthesis every two or three years.There are approximately 400,000lowerA prosthetic limb for a transtibial amputee consists of several components (Fig. 1). First there is a socket that fits over the residual limb that is custom fit for each amputee. A fitting is505attached to the lower end of the socket to attach a pylon.

1). First there is a socket that fits over the residual limb that is custom fit for each amputee. A fitting is505attached to the lower end of the socket to attach a pylon. The pylon connects to a prosthetic foot and is sized to the proper length for the prosthetic limb.The most important aspect of a lower extremity prosthesis is socket design. The socket is the interface between the human and the mechanical support system. Ultimately, the design and fit of the socket is what determines patient acceptance, comfort, suspension, and energy expenditure1. All of these factors in unison determine the real utility of the final product.The traditional method for designing sockets requires a skilled prosthetist and time consuming manual steps2, 3. First, the patient's residual limb is wrapped with plaster to acquire the shape. The plaster wrap is used to create a plaster positive.

First, the patient's residual limb is wrapped with plaster to acquire the shape. The plaster wrap is used to create a plaster positive. The prosthetist empirically modifies the plaster positive of the residual limb to distribute weight bearing to tissues that can tolerate the required forces. The socket is molded over the biomechanically corrected pattern. It is usually necessary to destroy the pattern in order to remove it from the socket. Making a new socket involves repeating all the steps if the pattern is destroyed.CAD and Computer Assisted Manufacture (CAM) techniques are beginning to be widely used to design and manufacture sockets. A mechanical digitizer, magnetic digitizer, or a non- contact laser scanner4 inputs the residual limb shape into the computer. The prosthetist then uses specialized software to produce a biomechanically correct socket from the limb shape. A computer controlled milling machine carves the pattern for the socket from plaster or foam.

A computer controlled milling machine carves the pattern for the socket from plaster or foam. The socket is then made using conventional methods such as vacuum molding or lamination.A limitation of currently available prosthetic CAD systems is that they only design the shape of the inner wall of the socket, which is all that is necessary for conventional fabrication. These CAD systems do not actually design the three dimensional shape of the socket. The actual wall thickness of the socket and the means of attaching the remainder of the prosthetic limb are determined during manufacturing. This is still largely an artisan process.SFF techniques are a good match for use in prosthetics. Several artisan steps can be eliminated. These include trimming the final molded socket and the addition of a pylon attachment fitting. Additional features can be included such as variable compliance socket walls and integrated fittings.

These include trimming the final molded socket and the addition of a pylon attachment fitting. Additional features can be included such as variable compliance socket walls and integrated fittings. Due to the nature of the manufacturing process, an increase in socket sophistication does not lead to increased cost. The SFF industry is largely service bureau based which is the mode in which many prosthetic facilities operate.Previous WorkOver the last ten years there have been several attempts to use SFF in the manufacture ofprosthetic sockets for amputees.Northwestern University working with Baxter Healthcare made a single transtibial socketusing Stereolithography (SLA) in 19905.In 1991, The University of Texas at Austin (UTA) and The University of Texas Health Science Center at San Antonio (UTHSCSA) collaborated to make scaled down transtibial506sockets6, 7. A full sized socket that incorporated a fitting for attaching the pylon followed this in 1992.

A full sized socket that incorporated a fitting for attaching the pylon followed this in 1992. An amputee in a supervised setting wore this socket briefly.In 1992 Rovick from Northwestern developed a rapid prototyping technology called Squirt Shape to address the need for rapidly fabricating prosthetic sockets8. The process is a form of Fused Deposition Modeling where a bead of molten plastic is extruded in a continuous spiral to form a single wall socket.In 1998 Freeman and Wontorcik fabricated check sockets suitable for ambulation though they lacked the durability needed for extended use9. The cost of the sockets was deemed to high for use as check sockets.Also in 1998, Lee et al reported on fabricating two prosthetic sockets for amputees using FDM10. Gait analysis was performed comparing conventional and FDM sockets.

Gait analysis was performed comparing conventional and FDM sockets. Minimal variations in gait between the two types of sockets were shown.In 1999 UTA and UTHSCSA collaborated on a sophisticated double wall socket fabricatedusing SLS11, 12. This work led to the current clinical pilot study.the UTHSCSA/UTA collaboration used SFF for anything except replacing the first manual step in socket fabrication. The remainder of the prosthetic limb had to be attached by artisan means and no advanced features such as variably compliant walls were incorporated.Itis importantto note that none ofthe above work with the exception ofMethodsThe previous work involving the double wall socket was extremely complicated and was burdened by the extra weight of the double wall configuration. It was decided to use a simpler approach for this pilot study using a single wall configuration with variable wall thickness.

It was decided to use a simpler approach for this pilot study using a single wall configuration with variable wall thickness. The approach was to vary the wall thickness to provide thinner more flexible socket walls over sensitive areas of the residual limb.The clinical feasibility of using SLS for socket fabrication was evaluated by comparing conventional and SLS sockets for comfort, stump-socket interface pressures and gait function in four transtibial amputees. Three were vascular amputees and one was a traumatic amputee. Their ages ranged from 31-62 and all would be considered community ambulators. Informed consent was obtained following local institutional review board protocols.A conventional patella tendon bearing socket was designed for each subject using standard CAD/CAM fabrication techniques. The residual limb of the amputee was cast at the UTHSCSA using a ShapeMate13 casting sock. This sock hardens in minutes after being put in water.

The residual limb of the amputee was cast at the UTHSCSA using a ShapeMate13 casting sock. This sock hardens in minutes after being put in water. A laser imager was then used to measure the outside shape of the ShapeMate sock. The sock has a uniform thickness so that by measuring the outside of the sock then subtracting the sock thickness, the correct residual limb shape can be calculated. The socket was designed using ShapeMaker13 prosthetic CAD software. A socket pattern was then milled from urethane foam. The conventional socket was fabricated using carbon fiber lamination. A pylon and prosthetic507foot were added to complete the prosthesis. Final alignment of the prosthesis was performed using visual gait analysis and patient feedback.Each subject was given at least two weeks to acclimate to the conventional socket before undergoing gait analysis. Reflective markers were placed on the sacrum, bilateral ASIS, knee, ankle, and 2nd metatarsal phalangeal heads.

Reflective markers were placed on the sacrum, bilateral ASIS, knee, ankle, and 2nd metatarsal phalangeal heads. Ground reaction forces were obtained using AMTI14 force plates embedded into the gait lab floor. Combined joint kinematic and kinetic data was collected at a sampling rate of 60 fps for three representative strides as the subject walked at his self-selected speed. Using VICON Clinical Manager software15, basic temporal and spatial features of the gait patterns and lower extremity joint angles were determined. Using inverse dynamics, net joint moment (torque) was calculated. For the prosthetic limb hip and knee joints this represented the net muscular generated joint moment that is developed in response to external forces and loads.

For the prosthetic limb hip and knee joints this represented the net muscular generated joint moment that is developed in response to external forces and loads. For the prosthetic ankle, the joint moment is the resultant internal moment developed at the foot/pylon junction that mirrors the location of the intact limb ankle joint.At the time of the gait analysis each subject filled a questionnaire regarding limb use,wearing schedule and rated socket comfort using a visual analog scale.The urethane foam pattern used to fabricate the definitive socket was measured using the laser imager and served as the template for SLS socket fabrication. This was to insure that the comparison was between sockets of identical shape. Trim lines, load bearing areas, and pressure sensitive areas were marked in black on the pattern. The laser imager captured these marks. The data including the marked locations were saved in an AAOP data interchange file.

The laser imager captured these marks. The data including the marked locations were saved in an AAOP data interchange file. This is a non- proprietary file format that is supported by prosthetic CAD/CAM vendors.In previous efforts at UTHSCSA/UTA9, 10, the process of taking an AAOP data interchange file and creating a complete 3D socket with an attachment fitting time consuming and intensive, was extremely labor involved several different software packages. In order to produce sockets in a timely manner it was necessary to write custom software to create the final STL file from the AAOP file. This software was written in C++ using the OpenGL graphics API for the Microsoft Windows platform. The software performs four main functions. First it gives thickness to the socket by creating an outer shell from the inner socket shape. The nominal socket wall thickness was set to 6 mm. The socket trimline is set interactively using the mouse cursor to select points on the socket shape (Fig. 2).

The nominal socket wall thickness was set to 6 mm. The socket trimline is set interactively using the mouse cursor to select points on the socket shape (Fig. 2). Areas of variable wall thickness are interactively drawn on the socket shape. These areas are defined relative to pressure sensitive areas marked on the socket pattern. The minimum thickness of regions over sensitive areas was set to 1.3 mm. Finally a fitting for a pylon adapter is added to the bottom of the socket. The fitting has a key that matches a keyway in the pylon adapter. The completed design is saved in an STL file (Fig 3).508emailed to Accelerated Technologies Inc.16, a service bureau, for fabrication. The parts were built using Duraform on a Sinterstation 250017. The sockets were coated with cyanoacrylate for a better surface finish after sanding.

The parts were built using Duraform on a Sinterstation 250017. The sockets were coated with cyanoacrylate for a better surface finish after sanding. The finished sockets were sent back to UTHSCSA for final assembly.Each STL files wasEpoxy was used to glue the nylon pylon adapter into the keyed fitting on the socket even though there was already a snug fit. The pylon and a prosthetic foot were added to complete the prosthetic limb (Fig. 4). The pylon adapter has two eccentric cylinders that are used to set the proper alignment angle of the pylon relative to the socket.The gait analysis and subjective evaluation In was repeated using the SLS prosthesis. addition, measurements were made using the FSCAN18 system to compare stump-socket interface pressures in the conventional socket and the SLS socket.

addition, measurements were made using the FSCAN18 system to compare stump-socket interface pressures in the conventional socket and the SLS socket. The FSCAN system uses pezioresistive pressure sensitive plastic strips that were attached to the skin of the residual limb at sites that corresponded to the thinned Interfaces the SLS socket wall. regions of pressures were recorded and averaged over 10 strides. Pressure measurements were first conducted while wearing the conventional socket and then repeated with the SLS socket without changing the location of the sensors.A finite element model of the socket was developed at UTA during the same time period though it was not ready to use during the socket design phase.

Using IDEAS from SDRC19 the socket volume is meshed with solid elements (parabolic tetrahedra), generating the final FE model that can be associated with selected material properties, undergo the appropriate boundary conditions and be simulated to generate structural analysis results. This finite element model proved to be valuable during the clinical part of the study.Materials’ testing was also done at UTA. Sample Duraform coupons fabricated at UTA were subject to a uniform pressure and the deflection was measured. It was necessary to verify that the material properties of the parts produced by SLS were the same as used in the finite element model.509ResultsA clinically acceptable fit was achieved in all subjects with both sockets. All subjects were satisfied with the fit of the SLS socket and expressed a willing to use it to meet their ambulation needs.

All subjects were satisfied with the fit of the SLS socket and expressed a willing to use it to meet their ambulation needs. Informal feedback from the subjects suggested that overall socket function was nearly identical in three subjects who had a history of satisfactory socket fit and prosthetic use. The fourth subject had a long history of stump pain and poor tolerance to prolonged prosthetic use. This subject displayed the greatest degree of improvement in comfort and use of his prosthesis when using the SLS socket. Using the visual analog scale (0-10 range) comparison of socket comfort between the sockets ranged from the same to 15% improved with the SLS socket. Average daily wearing time remained the same in three subjects and increased by 5 hours in the fourth subject.Gait analysis showed no significant differences between SLS and conventional sockets inself selected walking speeds or knee joint torques.Attempts at measuring interface pressures were problematic.

The difficulties were related to sensor movement between test sessions and bucking and wrinkling of the sensors during the loading/unloading cycle associated with each step. The resulting measurements contained spurious errors and the data needs to be interpreted cautiously. Considerable variability in the data existed but some trends were observed. Average peak pressures during the stance phase of gait showed no differences in the lateral aspects of the socket but were decreased by 11 psi in the anterior aspects of the SLS socket compared to the conventional socket.There was one notable failure where a socket fractured as one of the amputees stepped off of a bus. The socket subsequently broke in two pieces as the amputee continued to walk with the limb (Fig. 5). While this caused great concern about the whole project, it led to greater understanding of the whole process.

5). While this caused great concern about the whole project, it led to greater understanding of the whole process. Upon close examination it was revealed that there were both design and materials components to the failure.The materials problem resulted from the way the Sinterstation 2500 was set up. It was previously noticed that the socket material was quite brittle. This was the result of a tradeoff between feature definition and part density. The traditional way of operating an SLS machine is to use a laser power that is less than what would produce the densest possible part. This gives better feature definition and the part is easier to remove from the powder bed. Running the machine at a higher laser power results in a denser part but with a small loss in accuracy. The denser parts are stronger and have material properties that are closer to an extruded part. The socket that fractured was not fabricated at the highest possible density.

The denser parts are stronger and have material properties that are closer to an extruded part. The socket that fractured was not fabricated at the highest possible density. All succeeding parts were fabricated at a higher laser power and there were no further failures.510The design problem came from the the maximum desire compliance at the socket. This resulted in a notch between the pylon adapter and the socket. This was the region where the fracture started. The socket was redesigned to blend the socket and pylon smoothly together.toprovidethe distal end ofThe finite element model of the failed socket showed that at the point of failure was the location of maximum stress (Fig 6 in red). The revised socket showed reduced stress at the junction of the pylon and socket.

The revised socket showed reduced stress at the junction of the pylon and socket. In neither case did the maximum stress in the model approach the (theoretical) ultimate limit of Duraform.DiscussionThis study showed that the SLS sockets were clinically acceptable and that they can befabricated in a timely manner.The desired amount of compliance hoped for was not achieved by thinning the socket wall. It was judged that a minimum wall thickness of 1.3 mm was necessary for structural integrity and even at that thickness there was little flexibility.The pylon adapter chosen for this study presents an obstacle to clinical acceptance. The eccentric cylinder pylon adapter was chosen as it was simple to incorporate into the design. To be accepted by the prosthetics profession it will be necessary to accommodate conventional pyramid adapters.Future WorkLarger scale clinical studies are planned.

To be accepted by the prosthetics profession it will be necessary to accommodate conventional pyramid adapters.Future WorkLarger scale clinical studies are planned. In the future, sockets will be designed directly from scan data eliminating the pattern fabrication step. One important addition will be incorporation of industry standard hardware into the socket design. New methods for providing additional compliance for sensitive areas of the residual limb will be employed. It is planned to use finite element modeling of the sockets during the design phase to minimize the possibility of failure as well as testing new methods of varying the socket compliance.AcknowledgementsThe authors wish to acknowledge the Veterans Administration for their support of this study by the VA Research and Development Service Merit Review grant #A1963PA. The authors also wish to thank The DTM Corporation and Accelerated Technologies Inc.

OSED LOOP CONTROL OF 3D LASER CLADDING BASED ON INFRARED SENSINGD. Hu, H. Mei, G. Tao, R. KovacevicResearch Center for Advanced Manufacturing, Southern Methodist University 1500 International Parkway, Suite 100, Richardson, TX, 75081AbstractIn this paper, a heat input closed-loop control system based on infrared image sensing for 3D laser cladding is introduced. A high frame-rate (up to 800 frames/s) camera is installed coaxially on the top of the laser-nozzle setup. Complete of the infrared images of the molten pool can be acquired with a short nozzle-substrate distance in different scanning directions, eliminating the noise from the metal powder. The characteristics of the images show a clear relationship with the parameter variations of the cladding process. A closed-loop control system is built based on the feedback of the infrared image sensing.

A closed-loop control system is built based on the feedback of the infrared image sensing. The control results show a great improvement in the geometrical accuracy of the part being built.Keywords: SFF, Laser Cladding, Infrared Image Sensing, Control.1. IntroductionSolid Freeform Fabrication (SFF) is a novel manufacturing technology that could be used for the rapid creation of models, prototypes and patterns and for limited-run production. Nowadays, several SFF methods have been developed, such as 3D welding [1,2], micro-casting [3], selected laser sintering [4], Laser Engineered Net Shaping (LENS) [5], Shape Deposit Manufacturing (SDM) [6], Directed Light Fabrication (DLF) [7], 3D laser cladding [8] and some hybrid methods [9-11]. Among these SFF methods, 3D laser cladding has been considered to be very suitable for building metallic parts.

Among these SFF methods, 3D laser cladding has been considered to be very suitable for building metallic parts. A laser beam could be easily delivered and controlled and could be focused on a very small area, so that the interaction zone of a laser process can be accurately positioned and the bead track could have a small width (< 1 mm). This means that higher geometrical accuracy and surface quality of the final part can be achieved. Moreover, because of the powder delivery feature in 3D laser cladding, no inert-gas protection chamber is required. So, a larger part can be produced, and a more complicated cladding path can be traced, including four-dimensional and five-dimensional paths. By controlling the mass flow rate of the metal powders from different powder feeders, a composite material or alloy with a functionally gradient distribution can also be produced.129When laser cladding is used for SFF, it builds parts by adding powder material (e.g.

layer by layer) and in a way that it resolves a complicated manufacturing process into a series of simple and repetitive layer procedures. A stable and repeatable layer procedure is therefore crucial for the quality of the part produced. However, the laser cladding process is governed by a large number of parameters [11]. Some of the parameters are sensitive to the environment variations and some of them affect the other laser cladding parameters. In order to perform successful 3D laser cladding operations, an on-line closed-loop control of the process is necessary.Several previous research works [12-18] have been done to control the laser cladding process. Some of them are focused on controlling the powder delivery in order to achieve more stable powder delivery [12,13], or study the powder stream distribution and laser-powder interaction [14,15].

Some of them are focused on controlling the powder delivery in order to achieve more stable powder delivery [12,13], or study the powder stream distribution and laser-powder interaction [14,15]. Others deal with the optimization of the laser cladding process by a closed- loop control, using a CCD camera [16-18] or a phototransistor [19] as the sensing device. However, the CCD camera or the phototransistor is installed usually from one side of the nozzle. This installation is not practical for SFF 3D cladding. The distance between the nozzle and the substrate is very small (~5mm), so the viewing area of the cladding is limited. The images are deformed due to the angle between the optical central line of the camera and the laser-nozzle setup. Also, the images acquired from an asymmetrically installed camera will vary according to the scanning direction, which is always changing in SFF processes.This paper discusses the control of the heat input in 3D laser cladding for SFF.

A closed-loop control system based on infrared image sensing is developed. A high frame-rate (up to 800 frame/s) camera is installed coaxially on the top of the laser-nozzle setup. Complete infrared image of the molten pool and surrounding area can be easily acquired with a short nozzle- substrate distance and different scanning directions, eliminating the noise from the metal powder. The characteristics of the images vary with the cladding process parameters and have a clear relationship with variations of those parameters. A series of laser cladding experiments were conducted with and without the closed-loop control. The experimental results show a great improvement in the geometrical accuracy of the part being built.2. Development of closed-loop control systemThe 3D laser cladding system is comprised of four subsystems: a 1 kW Nd:YAG laser source, a powder delivery system, a 3-axis positioning system and an infrared image acquisition system.

The complete system is shown in Figure 1. The 3D metal part is produced layer by layer by synchronizing the X-Y motion, the Nd:YAG laser and the injection of metal powder. The infrared image acquisition system takes the images of the molten pool in real time and calculates the dimensions of the molten pool. It compares the recorded information to the preset value and creates a feedback control value to modify the output of the laser source. In this way, a stable molten pool with desired dimensions is obtained, and thus it will be possible to produce stable and repetitive layers if stable powder delivery is ensured.130The difficulties of image sensing in 3D laser cladding arise from several problems. The distance between the cladding nozzle and the substrate is short (~5 mm). This distance blocks the sight if the observation is from the sides of the nozzle.

The distance between the cladding nozzle and the substrate is short (~5 mm). This distance blocks the sight if the observation is from the sides of the nozzle. In order to observe the molten pool of cladding from one side, the camera is usually installed at a large angle with respect to the central line of the nozzle setup. The image acquired will be distorted due to that angle as well as to the direction of the motion. This problem will be exaggerated when the scanning path is multi- dimensional where the part could come in collision with the camera or could obstruct the direction of observation. Other problems of image sensing come from the intense light produced by laser cladding.

Other problems of image sensing come from the intense light produced by laser cladding. The metal powder adds to the difficulty of image processing as well.CCD CameraLaser headPowder Feeder3-axis Motion System1000W Nd:YAG LaserLaser ControlFrame GrabberMotion ControlLaser cladding Control CenterFigure 1 System Integration of 3D laser cladding.A coaxially installed image sensing setup with respect to the laser beam is developed to solve the problems mentioned above. The optical part of the laser head (Figure 2(a)) provides another optical path for the observation of the laser processing. A high frame-rate (up to 800 frames/s) camera is installed on top of the laser head, taking gray-scale images with a size of 128×128 pixels. The radiation from the molten pool produced by the laser cladding passes through the partially reflective mirror and an infrared filter, forming an image on the CCD chip of the camera.

The radiation from the molten pool produced by the laser cladding passes through the partially reflective mirror and an infrared filter, forming an image on the CCD chip of the camera. A Nd: YAG laser blocking filter (∼1.06 µm) is utilized to protect the camera from laser damage. An iris is used to adjust the intensity of the radiation received by the camera to prevent over-exposure. In order to get the infrared image of the molten pool, which reduces the high intensity light from the molten pool and eliminates the image noise from the metal powder, an infrared filter is selected (>700 nm) and installed between the iris and camera. During the laser cladding process, the camera acquires images of the cladding process in a frame rate of 200 frames/s. Images are transferred to the frame grabber installed on a PC computer that carries out the sensing and control process.

Images are transferred to the frame grabber installed on a PC computer that carries out the sensing and control process. The real system setup is shown in Figure 2(b).Figure 3 provides an infrared image acquired by the coaxially-installed camera under practical laser cladding condition. Because the observation is directly from the top of the molten pool, a full view image of cladding can be acquired without any blocking. With the correct combination of the Nd: YAG filter and the IR filter as well as the right iris aperture, a clear131image of the molten pool can be obtained without the noise from the metal powder. The radiation wavelength received by the camera is between 0.7 and 1.06 µm.Meriaudeau and Truchetet [16] calibrated the CCD camera with an IR filter, and concluded that the gray level of the pixels has a linear relationship with the temperature. So, the infrared images acquired reflected the temperature distribution around the molten pool.

So, the infrared images acquired reflected the temperature distribution around the molten pool. Due to the temperature range in laser cladding, the wavelength region of 0.7-1.06 µm occupies the part of the electromagnetic spectrum in which most radiometric surface temperature measurements are made. Figure 3(b) shows the gray level isotherms of the unprocessed infrared image shown in Figure 3(a). In a closed loop control of 3D laser cladding process, an absolute temperature measurement is not necessary, as only the relative difference is necessary for generating the feedback.High Frame-rate CameraFrame GrabberIR FilterNd: YAG LaserIris Nd: YAG FilterPartially Reflective MirrorPowder Delivery NozzlePowderPIII 350 ComputerSubstrateV(a) Schematic presentation(b) Close-upFigure 2 Infrared image acquisition system(a) unprocessed(b) processedFigure 3 Acquired infrared image1323.

Laser cladding experimentsA set of laser cladding experiments have been performed to test the control effect of the developed closed-loop control system based on infrared image sensing. The substrate material utilized in the experiments is mild steel, and the metal powder is an H13 tool steel powder with a -100/+325 mesh size. The powder delivery rate is kept constant at 5.2 g/min. The experiments are carried out at a constant scanning speed of 8 mm/s, but at different laser powers.To compare the processing results, single-bead walls are built under the following three varied processing conditions. The first type of samples are built with no pre-heating of the substrate and no closed-loop control for the heat input. The second type of samples are built with pre-heating but no control, and the third type of walls are built utilizing the closed-loop control function. The walls are produced in a repetitive manner.

The second type of samples are built with pre-heating but no control, and the third type of walls are built utilizing the closed-loop control function. The walls are produced in a repetitive manner. The nozzle moves along the positive direction of the X axis first to build one layer, moves up a small increment along the Z axis after the layer is built up, and then moves back along the negative X direction for the next layer’s deposition. During the wall building process, the area of the thermal field with a threshold of 50 is recorded continuously with an image-processing rate of 10 frames/second.A typical group of single-bead-wall samples built by laser cladding at the same laser power (290 W) is shown in Figure 4. It can be observed that, when no heat input control is utilized, the root of the wall is obviously narrower than the upper part of the wall, as shown in Figure 4(a) or (b).

It can be observed that, when no heat input control is utilized, the root of the wall is obviously narrower than the upper part of the wall, as shown in Figure 4(a) or (b). Moreover, this trend is even more pronounced within the wall built without preheating (Figure 4(a)). The geometrical change of the uncontrolled 3D laser cladding process results from the variation in heat loss. At the beginning of the wall building, due to the large heat conduction of the substrate, the created beads are narrower. As the wall grows up, there is less heat conduction along the built wall, and the bead becomes wider until it reaches a new equilibrium value. For the wall with preheating, the initial heat loss conducted into the substrate is reduced because of the high temperature of the substrate, and thus more energy is used to create the molten pool, which results in a wider wall than that without preheating.

The role of the closed- loop control system in laser cladding is just to overcome the effect of heat variation. At the initial stage of wall building, because a lot of heat is conducted into the substrate, a small molten pool is created. The control system reacts by increasing the laser power in order to produce a molten pool with the preset dimensions. As the wall grows, heat conduction loss is reduced and the laser power output returns to a normal level. With this feedback control function, the size of the molten pool could be kept nearly constant, resulting in the constant wall width along the entire height of the wall.This process can be more clearly explained with the use of the recorded infrared images recorded during laser cladding. As was mentioned before, the pixel number of the infrared image reflects the temperature distribution of the laser cladding as well as the size of molten pool.

As was mentioned before, the pixel number of the infrared image reflects the temperature distribution of the laser cladding as well as the size of molten pool. It can be observed from Figure 5(a) that, when neither preheating or heat input control is applied, the pixel number of the thermal field area starts with about 1,000 for the first layer, and it gradually increases up to 5,000 and stays there for the rest of the buildups. However, when a preheating procedure is applied, the pixel number of the infrared images at the beginning is relatively high compared to the case without preheating, as shown in Figure 5(b). This fact corresponds to the133geometrical comparison of the walls built under these two conditions. Figure 5(c) shows the pixel numbers of the images acquired from the laser cladding of a single-bead wall with heat input control. Apparently, a very stable pixel number is obtained which corresponds to a uniform width of the wall.

The average width of the wall wa is calculated by averaging 10 width measurements taken at 10 evenly spaced heights from the substrate surface. wmin means the minimum width of the wall, which normally corresponds to the root of the wall.

wmin means the minimum width of the wall, which normally corresponds to the root of the wall. A geometrical accuracy η is introduced to express the width variation of each wall sample, which is calculated by the following formula:η=− ww w aamin ×%100Table 1: Geometry of the walls built under different conditionsPower Level (W)290 370 450 540 620Without preheating and without control Minimum width wmin (mm) 0.45* 0.58 0.79 0.945 0.99Average width wa (mm) 0.95 1.33 1.48 1.61 1.7Geometrical accuracy η (%)53% 56% 47% 41% 42%With preheating and without control Minimum width wmin (mm) 0.62 0.9 0.94 1.0 1Average width wa (mm) 0.95 1.26 1.5 1.60 1.75Geometrical accuracy η (%)35% 29% 37% 38% 43%Average width wa (mm) 0.91 1.02 1.35 1.58 1.67With control **Minimum width wmin (mm) 0.90 1.01 1.15 1.29 1.31: The bonding of the wall with the substrate surface is not continuous.

**: The laser powers listed are transformed from the calibrated pixel numbersThe geometrical change of the walls with laser power is shown in Figure 6. When no heat input control is applied, a preheating procedure contributes to improve the geometrical accuracy of the wall, especially when the laser power is less than 500W. With the increase of laser power, the influence of preheating becomes of less importance.

With the increase of laser power, the influence of preheating becomes of less importance. This is because a higher laser power could compensate for the difference in heat loss due to the different thermal conditions)%( n o i t a i r a v y r t e m o e G605040302010No preheating, no controlWith preheating, no control With control0 200300400500600700Laser power (W)Figure 6 Geometrical accuracy Vs laser power135Geometrical accuracy η (%)1% 1% 15% 19% 22%It can also be observed that, under all laser power levels, the geometrical variation of the cross sections of the walls built with heat input control is much lower than those without control. When the laser power is below 400W, there is almost no width variation along the wall height if a control function is applied. Even at higher laser power levels, the geometrical accuracy gradually goes down, but it is still far superior to that without any control function.4.

Even at higher laser power levels, the geometrical accuracy gradually goes down, but it is still far superior to that without any control function.4. ConclusionsInfrared image sensing is an efficient sensing method for 3D laser cladding for SFF. Clear infrared images of the molten pool generated by the laser beam can be acquired easily by a coaxial camera even in the case of a short nozzle-substrate distance and in different scanning directions. The infrared image reflects the temperature information that is a significant parameter affecting the cladding result. The closed-loop-controlled 3D laser cladding based on the infrared image sensing can overcome the effects of thermal variation and thus achieve a consistent processing quality of 3D laser cladding.5. AcknowledgementThis work was financially supported by THECB, Grant 003613-0022-1999, NSF, Grants No. DM1-9732848 and DM1-9809198, and the U.S. Department of Education, Grant No. P200A80806-98.6. References1.

11-16, New York.19. Koch, J., and Mazumder, J., 2000, “Apparatus and methods for monitoring and controlling multi-layer laser cladding,” U.S. Patent #6,122,564.20. Tikare, V., and Griffith, M., 1997, “Simulation of coarsening during laser engineered netshaping,” Proceeding of the Solid Freeform Fabrication Symposium, August 11-13, Austin, TX, pp.699-707.21. Sankaranarayanan, S., Guo, W., and Kar, A., 1998, “Characteristics of laser-fabricated metal structures,” Materials and Manufacturing Processes, Vol.13, No.4, pp.537-554.137.

viewed PaperCLOSED LOOP CONTROL UTILIZING IN-SITU PATTERN PRINTING AND READING FOR QUALITY LEVEL DETERMINATION IN ADDITIVE MANUFACTURINGAja Hartman, Lihua ZhaoHP Laboratories, Palo Alto, CA, USAAbstractHP’s Multi Jet Fusion (MJF) is a powder-based technology that selectively melts polymer powder, using a fusing agent, in a layer-by-layer fashion to create 3D parts. One of the challenges for wide adoption of additive manufacturing is the assurance of the print process and part consistency through a cost-effective and non-destructive fashion. Non-destructive part quality measurements can be achieved through a method of printing two-dimensional patterns at desired locations throughout the part. The readability of these patterns provides either the signals to actuate process changes during the print or information on part quality during and after printing. This method can also be used for covert part marking to provide design intellectual property security.1.

This method can also be used for covert part marking to provide design intellectual property security.1. IntroductionAdditive manufacturing (AM), the process of building a part layer-by-layer, is a key contributor to the industrial revolution, Industry 4.0, which is characterized by the combination of automation and information technology in advanced production systems [1][2]. 3D printers are an example of these advanced production systems, utilizing data to improve production effectiveness and efficiency through design optimization and process control [3]. Additive manufacturing has benefits over traditional manufacturing methods due to increased design complexity, small batch production, and short lead times, which has led to the additive manufacturing industry growing 33.5% to 9.795 billion in 2018 [4].However, AM scale-up from prototyping to mass production has been difficult due to in-process variability that can affect part quality [5].

All AM technologies have sources of process variability that are dependent on their specific parameters and materials utilized in their print process. There are many different additive manufacturing technologies and materials, Multi Jet Fusion (MJF) being one of the powder bed fusion (PBF) polymer printing archetypes. The MJF print process starts with a computer aided design (CAD) model that is sliced to form a stack of 2-dimensional (2D) images. The MJF machine then begins printing pre-print layers by spreading powder and warming them using overhead warming in combination with fusing infrared (IR) emitting lamps. After an adequate warming base has been formed, the first CAD model 2D geometry slice image is printed using a black IR absorbing fusing agent on the powder and melted using the fusing lamps. A new layer of powder is spread on top of the previous layer and the process (print, fuse, spread) is repeated until the full object is formed.

A new layer of powder is spread on top of the previous layer and the process (print, fuse, spread) is repeated until the full object is formed. This printing process describes the basic MJF production system HP Jet Fusion 4200/5200 series 3D printing solution. However, MJF technology is not limited to utilizing only one fusing agent; it has the potential to perform in-situ chemistry on a voxel level by selectively jetting distinct functional agents.This MJF print process has several possible sources of process variability related to the different printing steps (print, fuse, spread). The MJF printer hardware functionality related to each of these steps can change due to wear. Image quality can suffer from these worn parts that are not detectable using current software and pen alignment tests. This variation can cause MJF parts to have different material properties and levels of quality.

This variation can cause MJF parts to have different material properties and levels of quality. Currently, detection methods for AM-produced part material properties require testing with a tensile tester, impact tester,1369hardness tester, or other mechanical equipment that permanently damages the part. There are also several non-destructive approaches for measuring 3D printed parts without damage, including ultrasonic flaw detection and Archimedes principal density measurement. However, these methods cannot be implemented in-situ of a 3D printing process and require expensive equipment, measurement time, and an expert to conduct calculations. The density measurement process in particular wets the part, which may not be ideal for some materials or applications, and the ultrasonic flaw detection is limited by the process being unable to detect small defect sizes and incapable of reaching certain depths into the part [6].

An efficient and effective non- destructive method for measuring the MJF print processes, hardware, and material properties of printed parts is needed to enable an easily adoptable additive manufacturing solution.2. Materials and Methods 2.1 Powder materialAll experiments in this paper were conducted with HP 3D high reusability PA12 fromHP Inc, Palo Alto, CA USA. PA12 material properties can be found in table 2 [7].Table 2. HP 3D high reusability PA12 material properties.Powder melting point (DSC) Particle size Bulk density of powderValue 187 ℃ 369 ℉ 60 𝜇𝑚 0.425 𝑔/𝑐𝑚3 0.015 lb/𝑖𝑛3Method ASTM D3418ASTM D3451 ASTM D18952.2 Sample PreparationSamples were produced using an MJF experimental platform.

This printer is outfitted with a multi-agent print carriage, clear IR absorbing fusing agent, cyan agent, yellow agent, magenta agent, black fusing agent, fluorescent agent, IR warming lamps, two 650 W tungsten halogen IR fusing lamps mounted on either side of the print carriage, print bed, powder supply, and spreader.2.3 Measurement MethodsThe image capturing sensor utilized for measuring code readability was an 8-megapixel iSight camera. The software for reading the quick response (QR) codes was the “QR Scanner” application.A Mitutoyo 500-197-30 Digimatic 200 mm stainless steel digital caliper was used tomeasure the geometry of 3D printed parts.3. Results and DiscussionMachine-visible contrasting agents were used to print QR codes onto MJF printed parts. The agent contrast is important due to the test pattern requirement to be easily detectible under specific lighting sources.

The agent contrast is important due to the test pattern requirement to be easily detectible under specific lighting sources. Examples of visible spectrum contrasting and UV spectrum contrasting patterns printed are shown in figures 1, 2 and 3. The amount of contrast necessary between the test pattern and background is dependent on the image capturing sensor and pattern detector. An 8-megapixel iSight camera image capturing sensor was used in combination with the QR Scanner application under both natural illumination and a UV light source. Print pattern generation was adjusted by modulating the QR code size to control readability of printed patterns sensitivity to print processes quality. Figure 1 shows three different sized QR codes1370with the largest 1 cm x 1 cm code chosen due to its sensitivity to print quality and easy detectability when fused correctly.Figure 1.

Figure 1 shows three different sized QR codes1370with the largest 1 cm x 1 cm code chosen due to its sensitivity to print quality and easy detectability when fused correctly.Figure 1. Quick response codes printed with cyan agent on a yellow part under ambient lighting with varying levels of fusing and readability.The fusing conditions and thermal uniformity of a part can be discerned by utilizing the readability of printed codes. Three parts were printed utilizing varying amounts of fusing energy to create different density yellow parts with cyan codes, shown in figure 1. The far-left part (a) is over-fused and was unable to be read using our test setup due to the excessive flowing of the melted polymer. The middle part (b) has nonuniform fusing, which can be seen by the color gradient ranging from dark to light as you view the largest test code from top to bottom; however, it was still readable using our test setup.

The far-right part is under-fused and unreadable due to the under-fused code not having a detectible amount of contrast.The criticality of fusing conditions and uniformity is dependent on the part quality specifications designated by the machine-readable code generated and can be controlled by changing the test pattern complexity. Using the printing pattern geometry shown, the middle part was determined to be within quality specifications.Other than print fusing quality, print quality can also be determined utilizing a similar method. This pattern recognition process was utilized to detect hardware and software deficiencies during the MJF printing process on an MJF test bed that were undetectable using other methods.The MJF test bed passed pen alignment tests and the parts shown in figure 2 passed caliper geometric measurement assessment. However, the print quality deficiency was detected utilizing the pattern detection method described.

However, the print quality deficiency was detected utilizing the pattern detection method described. Failure diagnosis determined the failure sources to be a combination of controller misprinting the pattern by printing on both the forward and backward passes, instead of just the forward passes, and backlash from belt wear, which caused the second pass printing to be misaligned with the forward pass print. This combination of software and hardware issues made small defects in part quality, which were undetectable using traditional tests.1371Figure 2. Visually contrasting quick response codes (yellow code on magenta part, magenta code on yellow part, and florescent code on nonfluorescent black part) printed with readable QR code on top and poor image quality unreadable code on bottom.Besides print and fuse quality determination, discreet embedded QR codes in 3D parts are also achievable.

An MJF fusing process and ink placement combined with a new functional agent that fluoresces under ultraviolet light sources enables this capability. To create the discrete fluorescent code, the following print process is utilized: one pre-fusing pass, one agent placement pass, and two fusing passes. For each powder layer, during the forward pass, there is one pre-fusing pass that heats the powder before the inks are printed to ensure splashing is minimized and good agent-powder interaction is achieved to promote the best image quality. This is accomplished by turning on the leading fusing lamp (mounted on the left-hand side of the printing carriage) during the forward pass. Directly after the leading fusing lamps (mounted on the right-hand side of the printing carriage) heat the powder, the ink is placed. Both leading and trailing fusing lamps are then turned on during the backward pass to fuse the layer.

Both leading and trailing fusing lamps are then turned on during the backward pass to fuse the layer. A new powder layer is spread, and the process is repeated.The MJF 3D printing experimental test platform has multi-agent capabilities with five of the agent channels utilized in these experiments. The agent setup and printing modes for the agent channels are shown in Table 1 for the printing of embedded fluorescent QR codes.Table 1. Agent channels and weight percent agent in nylon 12 polymer powder for producing discrete embedded fluorescent QR codes. Agent Channel1 Y 6.272 C 6.273 M 6.274 FA 0.01Agent Wt% Y – Yellow agent C – Cyan agent M – Magenta agent FA – Black fusing agent FL – Fluorescent agent13725 FL 17.60The fluorescent QR code dye overlaps with the printed part area dye, which makes it invisible under ambient room lighting conditions.

The fluorescent dye concentration of the QR code needs to be high enough to be readable by the sensor under ultraviolet light, yet low enough to be invisible to both the sensor and human detection during ambient lighting conditions. This phenomenon can be seen in Figure 3, where a poor unreadable image quality (lower) and readable image quality (upper) part are compared side by side under both ambient light and UV lighting conditions. The unreadable poor image quality part can either be used to discretely determine print quality or serve as a counterfeiting measure.Figure 3. Two printed gear parts with discreetly embedded fluorescent quick response codes under ambient lighting conditions (a) and ultraviolet lighting conditions (b).4. ConclusionA non-destructive method of printing two dimensional patterns at desired locations throughout the part with the readability of the pattern providing print quality information has been achieved.

This method enables covert part marking as well, to provide design intellectual property security to address the process-variability and intellectual property protection concerns of the manufacturing industry.A patent has been granted covering this method [8]. Although we have successfully demonstrated quality control using the readability of a printed pattern, a robust automated reading system implemented during the printing process and an automated pattern generation algorithm are being explored.5. AcknowledgementsThis work was funded by HP Laboratories. We thank Jim Stasiak and Greg Long for fluorescent agent support. The authors gratefully acknowledge the support of the HP 3D business extended 3D print team.13736. References [1] Schwab, Klaus. The Fourth Industrial Revolution. Portfolio, 2017.[2] Dilberoglu, Ugur M., et al. “The Role of Additive Manufacturing in the Era of Industry 4.0.” Procedia Manufacturing, vol. 11, 2017, pp.

viewed PaperCLOSED LOOP RECYCLING OF PETG IN FUSED GRANULE FABRICATION LARGE AREA ADDITIVE MANUFACTURINGN. R. Thompson* and J. M. Weaver*Department of Manufacturing Engineering, Brigham Young University, Provo, UT 84602AbstractPlastic waste is a critical worldwide problem that impacts additive manufacturing (AM). Extensive research has explored how plastic waste in AM can be reduced by recycling prints into new filament, with varying success. An alternative to filament-based extrusion is “fused granule fabrication” (FGF), which extrudes from pellets or granules. This method is often used for large area additive manufacturing (LAAM) of polymers. This paper expands upon the knowledge base from previous research on LAAM and examines the extent to which PETG can be recycled and reprinted through the same FGF tool without significant loss to its material properties. The metric used for comparing material properties is tensile testing along the direction of deposition.

The metric used for comparing material properties is tensile testing along the direction of deposition. Recycled material was granulized, filtered, and dehydrated. This resulted in effective printing of 100% recycled PETG, with recycled samples demonstrating 83% of the tensile strength of virgin PETG.IntroductionLarge area additive manufacturing (LAAM) is growing as a method for printing large structures [14]. The most common method of LAAM for polymers is fused granule fabrication (FGF) (also known as fused pellet fabrication or FPF). Although it does not have the precision or surface quality of traditional fused filament fabrication (FFF) processes, the speed at which the material is extruded is significantly higher [17]. FGF uses a feedstock of plastic particulates, usually pellets. The feedstock is fed through a screw-driven heating chamber and extruded through a nozzle. Aside from the appeal of printing quickly there is an added benefit to this process: machining [16].

The feedstock is fed through a screw-driven heating chamber and extruded through a nozzle. Aside from the appeal of printing quickly there is an added benefit to this process: machining [16]. Some FGF systems can be built to include a CNC milling tool, and some, like the AMBIT XTRUDE, can even be mounted in existing CNC machines [15]. The combined capability allows for hybrid manufacturing, attaining a greater surface finish than that which was originally printed [13]. Uses for this machine and its potential includes mold making, ability to use multiple materials, composite parts, and use in the automotive industry [1].These aspects of LAAM are appealing for higher volume manufacturing. Thanks to its rapid print speed as well as the inexpensive raw pelletized material, in comparison to rolls of filament, LAAM is becoming a desirable method for larger industries looking to make plastic parts swiftly [1].

Most polymer parts made commercially use virgin material because the cost for new materials is generally less expensive than reprocessing recycled material [18]. There has been a push for companies to “go green” and use a higher quantity of recycled materials in their plastic parts to help reduce both the quantity of plastic waste and the impact on our environment. Some estimates indicate that unless there is a change in how plastics are managed, the current trend of plastic waste will result in the weight of plastic in the oceans exceeding that of the fish that live in it by 2050 [9]. Utilizing additive manufacturing technology in industries can help reduce waste1100and offer sustainable value propositions for customers, but its effect is much more significant if fully recycled plastic can be used [11].This paper explores two important aspects regarding FGF and recycling.

First, the study considers the methodology needed to create granules from parts previously printed using the FGF process. Second, the study describes the process of printing with 100% recycled PETG using fused FGF methods and the resulting material properties. Previous research in FFF processes have been divided as to whether there is a significant difference in the tensile strength of fully recycled PETG in comparison to its virgin plastic counterpart. Some studies indicate that there is no difference [2], while others claim that there are significant differences, with further changes with subsequent cycles of recycling [7].

Some studies indicate that there is no difference [2], while others claim that there are significant differences, with further changes with subsequent cycles of recycling [7]. This paper examines whether there is a significant difference between virgin and full-recycled PETG for FGF processes by comparing material properties of the recycled plastic parts to virgin plastic parts in the same print orientation.Methods & ToolsPelletizing: The first steps taken were in determining a method to create particles small enough for the FGF tool head to use. Stock pellets are most often either cylindrical, or globular. There are other forms possible, but they are much less common [10].

Stock pellets are most often either cylindrical, or globular. There are other forms possible, but they are much less common [10]. We wanted to create a system that would be accessible to people with minimum equipment available, so we tried several common repurposed consumer tools in addition to an industrial granulator.We first attempted to use a countertop kitchen blender but found that the power of the electric motor was insufficient to slice plastic into consistent particle sizes. A paper shredder was also attempted, but it could only cut plastic around the thickness of 1.75mm filament before it was overwhelmed and jammed. A granulator machine (Figure 1) was able to cut the material easily, but the resulting particulates were too large and irregular to use in the FGF tool (Figure 2).Two methods could be used at this point to create particles of a compatible size. One common method is to remelt the particles and extrude them through a filament die.

One common method is to remelt the particles and extrude them through a filament die. Then the resulting filament can be sliced into more regularly shaped pellets [8][12]. However, this would negate any advantage of using an FGF tool instead of a filament system and would add an additional melting process between each cycle of recycling. Instead, we followed a second method, which is to feed the particles repeatedly through the granulator until they become more uniform. We found that after 6 to 10 cycles of running the parts through the granulator, most granules were the correct size.1101After the repeated granulator cycles, the particles were filtered and prepped for use in the FGF tool. For this, a double filtration device was created. The first screen has holes slightly larger than the maximum dimension of compatible pellets. This allows for the parts too large to be separated and granulated again.

The first screen has holes slightly larger than the maximum dimension of compatible pellets. This allows for the parts too large to be separated and granulated again. The second filter has holes slightly smaller than the minimum dimension of separate compatible particulates that would be too small to be used.pelletstoFigure 2. Industrial granulator.Figure 1. Recycled material after one granulator cycle (left) next to standardized pellets (right).Once the granules were sorted, they were placed in a dryer for 8 hours to help eliminate any moisture in the plastic. Printing without drying resulted in discoloration, void formation, and lower tensile strength [19]. Washing the plastic before drying was attempted, but we found no significant effect compared to simply filtering the plastic granules. After initial drying, the granules were placed in a drying cabinet for storage, then placed in a drying hopper prior to print.

After initial drying, the granules were placed in a drying cabinet for storage, then placed in a drying hopper prior to print. The plastic particulates sat in the drying hopper for 8 hours at 70 degrees Celsius before printing the material.Printing: Once the plastic granules were prepared, print extrusion tests of recycled PETG began. The granules were extruded on an AMBIT XTRUDE system mounted on a HAAS TM-2P machining system (Figure 3). Several test prints were done at the identical settings to that of the virgin PETG, but more consistent material flow and deposition occurred with spindle speed and print speed reduced by 10% and temperature raised by about 5 degrees compared to the settings for virgin PETG. All these adjustments were made to optimize the flow of the recycled plastic.

All these adjustments were made to optimize the flow of the recycled plastic. Through the slicing software, the part was printed in a continuous motion which slowly raised in the z axis, allowing for no seams in the print and a more homogonous appearing part.There was an occasional problem that arose where the flow of plastic ebbed due to the plastic granules no longer falling into the chamber where the screw is, so agitating the feed tube regularly was needed to counteract this problem. It is possible that the particulates are locking into place due to the jagged edges and corners found on the granules. This is potentially avoidable if further processing is done to smooth the corners, but no method was explored to discover a way to do so. Another solution would be to have an agitator of some form lightly jostling the material or tube wall, allowing the plastic to fall easily into the tool.Figure 3.

Another solution would be to have an agitator of some form lightly jostling the material or tube wall, allowing the plastic to fall easily into the tool.Figure 3. AMBIT XTRUDE FGF tool head in CNC mill.1102Parts for cutting tensile samples were printed as shown in Figure 4, where the long sides of the octagon are more than 165 mm in length to allow for variation of part placement in the vise when the machining process occurs. After print completion, they were examined to ensure that no foreign contaminates were found within the walls of the parts.Machining and Testing: The last step for testing was creating samples for machining the part. Cuts were made near the very end of the long sides of the samples, leaving as much material as possible (see Figure 5). These straight samples cut into samples by CNC.tensileFigure 4. Recycled PETG print.The cutting of the samples was done via two CNC programs to match Type 1 dimension specifications per ASTM D638.

Recycled PETG print.The cutting of the samples was done via two CNC programs to match Type 1 dimension specifications per ASTM D638. This was done using a two fluted half inch end mill, using climb cuts to leave the best finish on the part. Climb cuts were selected due to the superior surface finish in comparison to conventional cuts [5]. A modified vise was used to hold the plastic sample upright. This vise has a portion machined to hold the part in both its raw form and to hold the machined portion of the par when it is flipped over to have the part completed. Once the samples were completely machined, they were placed in the drying cabinet to ensure no moisture was introduced. Finally, parts were tested in tension on an Instrom machine and data was collected.Results and DiscussionThe recycled samples were found to have an average tensile strength of 56.00 MPa, with a standard deviation of 1.22 MPa (see Table 1).

The virgin plastic material, tested in the same fashion and with the same print orientation, the average tensile strength was found to be 60.06 MPa, exemplifying that there was a 6.76% decrease in overall strength for fully recycled PETG. However, as noted before, the recycled plastic required a 10% decrease in print speed and approximately a 5% increase in print temperature to achieve consistent, optimized flow. These changes must also be acknowledged when considering the difference in material properties.Figure 5. Cut lines for tensile samples.1103Table 1. Measured tensile properties.MaterialAve. Tensile StrengthSt. Dev.100% Recycled PETG56.00 MPa1.22 MPa100% Virgin PETG60.06 MPa0.67 MPaThese results were encouraging. Previous research [6] had demonstrated the strength of fully recycled printed parts to be approximately 60% of the original strength of virgin plastic.

Previous research [6] had demonstrated the strength of fully recycled printed parts to be approximately 60% of the original strength of virgin plastic. The more favorable results (93% strength) of this experiment may be due to the adjustments in print speed and temperature. It may also be due to the reduced processing of the plastic; the cited research remelted the plastic between prints to form the new feedstock. However, it should be noted that in a different study analyzing recycling of PETG, there was an increase in the tensile strength on the second and third cycle of recycling [7].Another possibility for the higher strength could be due to the nature of this manufacturing process. Other studies have been done using filament printers, which tend to have a smaller extrusion diameters and slower deposition rates.

Other studies have been done using filament printers, which tend to have a smaller extrusion diameters and slower deposition rates. It is possible that the larger extrusion pattern allows for better heat transfer between layers, allowing for the plastic between layers to bond to one another slightly better.ConclusionsAs shown in this paper, in-house closed-loop recycling of low friction plastics is possible, and it is a promising avenue for more sustainable LAAM manufacturing. The practicality of the system described here is not yet well optimized for labor and time. However, doing such a process allows for parts with 93% of the strength of virgin plastic in PETG, exceeding the strength of some other additive manufacturing processes.Potential opportunities of further research stemming from this can branch into a few different directions. Future research could examine the change of properties through successive recycling and reprinting of common polymers in FGF.

Future research could examine the change of properties through successive recycling and reprinting of common polymers in FGF. Another area of study is the effectiveness of different agitators in this process with jagged edged pellets/granules. One other avenue would be research in an improved method for the granulation of plastics that can be utilized immediately for FGF, injection molding, or other plastic-based manufacturing processes. A direct comparison of virgin PETG and fully recycled PETG could also be done using the printing parameters for the fully recycled material to allow for greater understanding of the effects of this recycling method and its direct translation to loss of material properties. Lastly, analysis of the strength between layers of 100% recycled plastic in FGF would also be beneficial to understanding the material’s properties.1104Citations[1] Crisp, Tyler G., and Jason M. Weaver.

An important aspect of PBF directly affected hereby is the coalescence behavior, as it significantly determines the surface quality, porosity, and thus the component’s mechanical properties. However, detailed studies on coalescence are limited to virgin powders with low viscosity; therefore, coalescence behavior of polyamide-12 with different molecular weights is investigated using hot stage microscopy. Additionally, the zero-shear viscosity is determined using the Carreau model, allowing comparison of experimental results to sintering models. Furthermore, the mechanical properties and surface qualities are analyzed, and components with adequate values are made with two-cycle reprocessed powder.

Furthermore, the mechanical properties and surface qualities are analyzed, and components with adequate values are made with two-cycle reprocessed powder. Since surface flaws do not exist uniformly across all components, the orange peel effect is not attributed solely to the increased viscosity of the reused powder.IntroductionThe term “Additive Manufacturing" refers to a wide range of technologies that all share a unique building method; it is a form of manufacturing where the three-dimensional (3D) product is built into its designed shape using a layer-by-layer approach directly from a computer-aided design (CAD) model [1]. This makes AM fundamentally different from conventional formative or subtractive methods, because it is a tool-free process, and it has several independent units creating a significantly larger degree of complexity in terms of cost-effectiveness.

In the beginning, a thin layer of powder, previously heated to the designated process temperature, is spread with a roller or blade onto the platform inside the building chamber all while being under an inert gas causing partial melting and compaction of the particles. The fusing of the melted particles has a critical impact on the porosity of the manufactured part, and thereby on the mechanical performance [2]. Since the process takes place under atmospheric pressure, the coalescence behavior plays a significant role here as the surface tension and viscosity of these particles is the driving factor. Most coalescence models are traced back to Frenkel. For Polyamide 12 in PBF, it is known that one of the predominant effects of process-induced ageing is post-condensation with linear chain growth yielding an increase in viscosity [3]. Although there are many studies on its influence of ageing and mechanics, the coalescence behavior of aged powders has never been directly evaluated.

Although there are many studies on its influence of ageing and mechanics, the coalescence behavior of aged powders has never been directly evaluated. Many studies only consider aged powder for qualitative statements like shifting the melting onset to higher temperatures and extending the coalescence time compared to virgin powder [4]. Furthermore, there are not many precise mechanical parameters for viscosities; rather, there are only generalizations about the process, such as the requirement that low viscosity is needed to achieve good coalescence with no porosity remaining in the fused regions. However, there are little investigations into the viscosities with which components are still produced and what properties they exhibit. The presumed "orange peel" surface flaw is said to be a result of high viscosities, so it is crucial to comprehend powder with increased molecular weight and investigate the surface to establish process quality rules.

The purpose of this paper is to describe the coalescence behavior of various aged PA12 powders with differing molecular weights, and their resulting mechanical performance and surface effects.40State of the ArtCoalescence is a phenomena by which at least two drops of a liquid or melt merge during contact to form a lager single particle [5]. The first studies were carried out by Frenkel [6]. Since the model is only valued at early stages of coalescence as the change of the particle radius is not considered, different authors adopted that model, and all determined that the properties of viscosity and surface tension are the determining factors.

The coalescence behavior of two particles is fitted with the modified Frenkel model for viscous sintering of liquids under isothermal conditions given by [7]:𝑑𝜃 𝑑𝑡=Γ 𝑟𝑜η2−5/3 cos 𝜃 sin 𝜃 (2 − cos 𝜃)1/3 (1 − cos 𝜃)(1 + cos 𝜃)1/3where Γ is the surface tension, η is the viscosity, a0 is the particle radius, θ is the angle between the line connecting the particle centers and that connecting a particle center with the extreme point of the neck.Figure 1: Modified Frenkel model according Pokluda et al. [7]Many authors have determined the behavior for materials and temperatures for PBF. Zhao et al. [8] showed that for virgin PA12, the Pokluda model describes the coalescence behavior sufficiently, and for measured surface tension, only a small effect of the temperature was observed. Haworth et al.

Haworth et al. [9] compared the viscosity of virgin PA12 powder with mixed powder (oven test 170 °C 5 h), and observed an increase in viscosity, and furthermore, calculated with the Frenkel model, the sintering kinetics. Another paper dealing with coalescence behavior for PA12 by Chatham et al. [10] focused on different process parameters and their influence on mechanical properties and density. For this purpose, the temperature in the process was observed with the aid of a thermal camera, and the necessary temperature-dependent characteristic values were taken from the literature to calculate the sintering process. Dadbakhsh [2] demonstrated that aged PA12 powder melts at slightly higher temperatures in hot stage microcopy for dynamic measurements with a heating rate of 10 K/min. Berreta et al. [11] investigated the necking time of high temperature materials PEEK and PEK using different MFI and hot stage microscopy processes.

Berreta et al. [11] investigated the necking time of high temperature materials PEEK and PEK using different MFI and hot stage microscopy processes. The authors determined that the lower molecular weight PEK fuses much faster (roughly by a factor of 2), but the tensile strength is reduced by 30%. According to Haeri [12], the aspect ratio of the parts has an impact on PEK. Plummer [13] investigated the effect of re-used elastomer powder and found that it had no effect on the hot stage microscopy or the mechanical behavior. Benedetti [14] was the first to describe different approaches for using diameter to calculate neck growth. More important, the use of different substrate plates for hot stage microcopy was determined, and only minor deviations were seen.

More important, the use of different substrate plates for hot stage microcopy was determined, and only minor deviations were seen. Furthermore, the diameter of the particle decreased as the temperature increased on a heating ramp.41ExperimentalIn this work, the common Polyamide 12 (PA2200, EOS GmbH) was investigated in three different stages, and processing tests were carried out without refreshing the powder. This technique resulted in a comparable temperature-time load in the used powder. Starting with virgin powder, three processing cycles were produced. In the first cycle, three construction jobs were fashioned, in the second, two construction jobs, and in the third cycle one construction job.

In the first cycle, three construction jobs were fashioned, in the second, two construction jobs, and in the third cycle one construction job. Each building job was 15.5 h; the according building time for each cycle is presented in Table 1.Table 1: Materials used Cycle NumberCount Build JobsStarting PowderRefreshingCumulative Building Time1 2 33 2 1Virgin powder Powder cycle 1 Powder cycle 2No No No15.5 h 31 h 46.5 hProcessing The processing was carried out on a Formiga P 110. Figure 1 (right) shows the construction job layout used. For the detection of local component properties, test specimens were manufactured in the xy and z directions. The resulting job height is 250 mm, and the filling level is 7.4 % by volume.

For the detection of local component properties, test specimens were manufactured in the xy and z directions. The resulting job height is 250 mm, and the filling level is 7.4 % by volume. Per each construction job, a total of 86 multi-purpose test specimens according to DIN EN ISO 527-1 type A were manufactured to determine mechanical properties under tensile load, and 91 80x4x10 mm³ test specimens were used to determine the component density with the following parameter settingsHeating time before start of job Resulting construction time Cooling time Exposure parameters Building bed temperature Builder chamber temperature Layer Thickness2.5 h 15.5 h 10 h EOS Standard 168 °C 150 °C 0.1 mmFigure 2: Process parameters (left) and Build job layout (right)Viscosity Number The viscosity number was determined by means of an Ubbelohde viscosimeter according to DIN EN ISO 307. Sulfuric acid was used as the solvent instead of the m-cresol specified in the standard.

Sulfuric acid was used as the solvent instead of the m-cresol specified in the standard. For each building job, three random samples were analyzed.Component Density The component density was determined using the weighing method on the components (see [15]). The component’s width and thickness were determined with a micrometer (± 0.001 mm) at three points, the component length was determined with a digital caliper (± 0.01 mm), and the component weight with a precision balance (± 0.001 g).Rheometer To understand the difference between the starting powders during the process, the zero-shear viscosity was determined by means of frequency sweeps in oscillation mode with a deformation γ of 0.005 via Carreau model at three different temperatures (200, 210 and 220 °C). The measurements were carried out on rational rheometer 42(HAAKE Mars 60) equipped with a 25 mm diameter; for quick and good thermal settings, the rheometer is equipped with a heated upper plate.

The measurements were carried out on rational rheometer 42(HAAKE Mars 60) equipped with a 25 mm diameter; for quick and good thermal settings, the rheometer is equipped with a heated upper plate. All measurements were conducted in the linear viscoelastic (upper limit γ 0.0045) range that was determined by amplitude sweeps. Immediately after post condensation takes place, conventional mold sample preparation is not possible. Instead, powder tablets were prepared with a press to guarantee consistent loadings and improve the consistency of the measurements. Furthermore, the samples were dried for one week in a vacuum oven at 70 °C. The complex viscosity η* was used to characterize the flow behavior as simple polymers are known to obey the Cox-Merz rule [16].Hot Stage Microscopy Hot microscopic observations for the sintering of polymer particles were performed using a Linkam FTIR600 heating and freezing microscope stages, controlled with central processor Linksys 32 software.

A glass plate was used as the substrate for the sintering particles. Two polyamide 12 particles from each cycle of approximately equal size were placed on the stage at ambient temperature and heated up to the isothermal temperatures of 200 °C and 210 °C with the max. heating rate of 150 °C. Preliminary tests ensured that the particles are homogeneously heated by the minimal wetting of the sample carrier with silicon lubricant [17]. The analysis of the experiments was performed using ImageJ software. Here, the sinter growth over time was evaluated by determining the dimensionless parameter of the sinter neck radius. Then the modified Frenkel model was solved in MATLAB by classical Runge-Kutta- method and compared with the experiments.ResultsPowder CharacteristicsViscosity Number The measured viscosity numbers (VN) of the new and used powders are shown in Figure 3.

It is seen that the viscosity number increases with each processing cycle, starting from VN = 60 ml/g for new powder to 136 ml/g after three processing cycles. With each processing cycle, the rate of change of VN decreases, and the viscosity number approaches a limit value. The increase in the viscosity number is attributed to post-condensation in the solid phase [18]. Due to the temperature-time stress during the process, the chain ends of the Polyamide 12 can react with each other, splitting off water. This leads to an increase in the molecular weight and a concomitant increase in the melt viscosity of the material.viscosity number in ml/gcum. building time in hbuilding height 250 mm, fill level 7.4 %01020304050406080100120140Figure 3:Viscocity number in dependence of cumulative building time (n=3)Hot Stage Microscopy In Figure 4, the isothermal hot stage measurements for the different cycles at a temperature of 200 °C are shown.

It is seen that the coalescence behavior is increasing significantly with increasing powder state. The fusing time for cycle 1 (virgin powder) is only a few seconds, whereas for cycle 2, the necking area forms after 10 s. For cycle 3, the necking area has only just formed after a time span of 30 seconds, and for complete fusing, almost 5 minutes was necessary. Dadbakhsha et al. [4] was also able to observe optical differences in the melted particles. It is assumed that the underlying cause may be added additives or optical artefacts.43C1C2C3Figure 4: Coalescence behavior at 200 °C for particles of different cyclesViscosity Measurements Figure 5 shows the different viscosity measurements conducted with the rheometer of the different powder states.

The loss and storage modulus dependence of the angular frequency of the different powders is depicted at 200 °C on left, the crossover point (COP) of these moduli is an indicator for the molecular structure of the samples as with increasing molecular weight the storage modulus increases and the COP moves to lower frequencies [19]. For cycle 1, no COP occurs at 200 °C, this is a result of the chains being so short that the storage modulus does not intersect that of the loss modulus even for the maximum possible instrument angular frequency of 628 rad/s. Because of the known post-condensation of the powder, the crossover point is detected for cycle 2, which shifts to even lower angular velocities for cycle 3. The zero-shear viscosity of the powders for 210 °C are displayed in the center. The measurements are in strong agreement if the increase of the different states is considered. With increasing cycle length, the increase in the zero-shear viscosity reduces.

The measurements are in strong agreement if the increase of the different states is considered. With increasing cycle length, the increase in the zero-shear viscosity reduces. It must be noted that the zero-shear viscosity cannot be analyzed by oscillation mode at temperatures lower than 210 °C, as the molecular weight of the powder is too high, and thereby the plateau is not reached as shown in Fig.

5 (right) for a temperature of 200 °C.storage modlus G´ in Paloss modulus G'' in Paangular frequency in rad/sG'' cycle 2100101102103104105106G'' cycle 3G' cycle 3G' cycle 2T = 200 °CCOPCycle 2Cycle 3Carreau fitCycle 110-1100101102100101102103104105complex viscocity in Pasangular frequency in rad/sT = 210 °C10-1100101102103104105complex viscocity in Pasangular frequency in rad/sCycle 3Cycle 2T = 200 °CFigure 5: Rheometer measurements, change cross-over at 200 °C (left), frequency sweeps at 210 °C (center) and 200°C (right) 44Coalescence Behavior Experiments The first isothermal temperature for evaluating the neck-growing radius at 200 °C is shown in Fig. 6 (left). As already illustrated in Fig.5, the progress of the neck radius shows extreme differences among the different powder states.

6 (left). As already illustrated in Fig.5, the progress of the neck radius shows extreme differences among the different powder states. Already seen in Cycle 1, sintering was observed at 196 °C; therefore, the measurement began already at a neck ratio of 0.2 However, the coalescence is completed after 7 seconds, which aligns with previous measurements of the virgin PA12 powder [20]. When considering the other two systems, the coalescence time to achieve a similar dimensionless radius is increased by one magnitude. For Cycle 3, there was a change detectable after 3 minutes, so a plateau was reached only after 240 s. For the second investigated temperature of 210 °C, cycle 1 was completely fused by reaching the holding temperature even with a high heating rate of 150 K/min. Therefore, the coalescence progress at 210 °C is only shown for cycle 2 and cycle 3.

Therefore, the coalescence progress at 210 °C is only shown for cycle 2 and cycle 3. In comparison to the first holding temperature, now cycle 2 and 3 are already partly melted and show already a necking during the heating up phase. The progress at the beginning shows a faster process, as expected, but the significant difference to 200 °C occurs towards the end, where the time to reach the plateau is halved. For the models, the previously determined viscosities were used as input variables and the surface tensions (35.6 mN/m) were used for new powders from the literature [21], and kept constant for all cycles. This simplification is permissible after the surface tension only has a weak function of temperature and molecuarl weight [22] in comparison to the the zero- shear viscoity. The resutls for the modfied Frenkel model at 210 °C are shown in Fig.

The resutls for the modfied Frenkel model at 210 °C are shown in Fig. 6 (right).Cycle 20601201802403000.00.20.40.60.81.01.2Cycle 1dimensionsless neck ratio a/rocoalesence tim in sCycle 3T = 200 °CCycle 20601201800.00.20.40.60.81.01.2Cycle 3dimensionsless neck ratio a/rocoalesence tim in sT = 210 °CCycle 3dimensionsless neck ratio a/rocoalesence time in sCycle 205101520Cycle 11,2 1,00.80.6T = 210 °C0.40.2Figure 6: Dimensionless neck ratio at 200 °C (left) and 210 °C, comparison with Pokluda modelIt is immediately noticeable that ratios around 1 and smaller are achieved in the experiments, whereas the models achieve ratios of 1.26; these deviations are known and discussed in the literature [23]. It is worth noting that for all experiments, irrespective of temperature or viscosity, progress is faster in the model than observed on the hot stage microscopy.

It is worth noting that for all experiments, irrespective of temperature or viscosity, progress is faster in the model than observed on the hot stage microscopy. Whereas the model predicts the neck ratio in a comparable time for cycle 1, but the differences increase for cycle 2 and 3 significantly.Part CharacteristicsPart Density Figure 7 shows the measured global component densities of the three processing cycles over the cumulative build time of the initial powder. A distinction is made here between specimens built in the xy and z directions. The component densities in both the xy and z directions are strongly dependent on the ageing state of the starting powder. The component density drops in the xy-direction from approx. 0.95 g/cm³ to 0.83 g/cm³. In the z- direction, this drop is more pronounced, with the component density falling from 0.92 g/cm³ to 0.72 g/cm³.

0.95 g/cm³ to 0.83 g/cm³. In the z- direction, this drop is more pronounced, with the component density falling from 0.92 g/cm³ to 0.72 g/cm³. Furthermore, the standard deviation of the component density increases due to a decreasing bulk density and an associated increase in the initial pore diameter, as well as an increase in surface roughness up to the "orange peel".45part density in g/cm³cum. building time in hxy-direction051015202530350.000.700.750.800.850.900.951.001.05051015202530350.000.700.750.800.850.900.951.001.05part density in g/cm³cum. building time in hz-directionFigure 7: Component density depending on the ageing status of the starting powderPart Surface An analysis of the parts in z- direction illustrates that the surface flaw “orange peel” occurs for cycle 2 at front and back position.

In the next cycle, positions of the parts with orange peel increases; however, even with a starting powder which has experienced 31 h of building time, parts without any surface flaws are produced.31 h (cylce 2)15,5 h (cylce 1)15,5 h (1 cylce) 31 h (2 cycle)Position “Orange peel“EOS PA220015,5 h (cylce 2) 31 h (cycle 3)Figure 8: Orange peel in dependence of part position and starting powderMechanical Performance There is a significant decrease of the elongation at break and the yield stress for z tension samples. Smaller decrease with in xy direction. In general z tension samples react more sensitively to aged powder. A substantial decrease of the elongation at break and the yield stress for z tension samples are observed, and the distribution of the elongation at break largely increases with cumulative construction time.

For cycle 3, the mechanics were not investigated after the density results dropped and increased surface defects occurred.Table 2: Mechanical CharacteristicsCycle nr.Cycle 1Cycle 2Build directionxy z xy zUTS in MPa51.22 (± 1.3) 48.35 (± 0.77) 47.96 (± 1.05) 22.26 (± 12.82)Elongation at break in % 15.78 (± 3.38) 11.28 (± 3.38) 15.16 (± 2.53) 4.24 (± 6.6)46Summary and OutlookThis study presents a theoretical and experimental investigation of the sintering process of commercial polyamide 12 powders as a function of the viscosity. When observing the viscosity number as well as the zero-shear viscosity according to Carreau, it is apparent that the ageing progress is maximum at the beginning and decreases with time seemingly reaching a plateau.

The results of the hot stage microscopy show that the growth of the sinter neck agrees with the modelling for the first cycle which is comparable to existing studies The coalescence behavior differs significantly, as cycle 1 powder investigations are not possible at 210 °C, since the particles are already coalesced despite a heating rate of 150 K/min. However, when the results are compared for powders with increasing molecular weights, viscosity deviations detected. The results achieved show a lower rate of coalescence than what the models discussed in this paper predicted. The following points are of importance when considering the manufactured components: the influence of cumulative construction time is more significant for components built in z direction. Additionally, the span of the resulting component density increases with increasing cycle time and is more distinct for z components. The mechanics achieve typical characteristic values for cycle 1.

Additionally, the span of the resulting component density increases with increasing cycle time and is more distinct for z components. The mechanics achieve typical characteristic values for cycle 1. For cycle 2, the UTS demonstrates a strong correlation to the building direction; where in the xy direction changes are small, and the z components only reach values of 50 % in comparison to cycle 1. The surface flaw "orange peel" shows a dependence on the position in the build platform; so, it is the first time for cycle 2 to occur in the rear and front area. However, despite further thermal stress around 15.5 h, and a significant increase in melt viscosity, components are possibly produced in the center without surface defects for cycle 3. This may be attributed to the different position dependent temperature profiles. The experimental coalescence studies confirm that the temperature has a significant influence on coalescence behavior.

This may be attributed to the different position dependent temperature profiles. The experimental coalescence studies confirm that the temperature has a significant influence on coalescence behavior. Here, scientific studies confirm that with adapted process parameters, processing is also possible with aged powder material. Regarding viscosity, it is important to note that a zero-shear viscosity of approx. 8*103 Pas for polyamide 12 at 25 K above melting point Tm ([4]) represents a limit for standard exposure. For further investigations, the coalescence behavior of mixed powders as it is typical in the process must be examined. Furthermore, the findings will be transferred to other material systems for the PBF-LB/P. To enable a general statement to be made about the material requirements in terms of viscosity.Acknowledgement Funded by the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation) – Project-ID 61375930– SFB 814, sub-project A3, T3.References1. 2.3.

ated Feedstock for Fabrication of Ceramic Parts by CAM-LEMZhien Liu, N. Suppakarn, and James D. Cawley Department of Materials Science and Engineering Case Western Reserve University 10900 Euclid Avenue Cleveland Ohio 44 106-7204AbstractIn laminated object manufacturing of ceramic components, lamination is one of the most important materials issues. Good lamination ensures monolithic component after firing. Otherwise, lamination defects that inevitably will occur in the parts will affect the properties of ceramic components. Adhesive (both liquid and non-liquid) lamination processes were developed for the cut-then-stack (CAM-LEM) procedure. The non-liquid adhesive lamination is discussed in detail.IntroductionCAM-LEM (Computer-Aided Manufacturing of Laminated Engineering Materials) is one solid freeform fabrication technique, based on cut-then-stack procedure, which uses ceramic green tape as feedstock.

One of the advantages of CAM-LEM is that ceramic parts with complex internal and external shapes can be readily produced. To achieve this requires that zero, or very small, strain is induced during stacking and lamination. One satisfactory method, called adhesive lamination, has been developed for CAM-LEM and reported elsewhere[l]. Generally, a mixture of a solvent, nonsolvent and diluent are used as adhesive. A variety of ceramic green tapes have been developed and tested [2-51, such as debased alumina, pure alumina, silicon nitride, PZT (lead-zirconate titanate) and ZTA (zircinia toughed alumina) tapes. By adjusting the details of the adhesive formulation and lamination process, 100% lamination efficiency and very low lamination strain (0-0.2%) were obtained for all systems.

By adjusting the details of the adhesive formulation and lamination process, 100% lamination efficiency and very low lamination strain (0-0.2%) were obtained for all systems. Recently, a non-liquid adhesive was also developed for CAM-LEM 161.Materials and experimentalThree kinds of ceramic green tapes were used in this work: a commercial (Coors') alumina tape, in-house alumina tape and in-house silicon nitride tape. Properties of these feedstocks are shown in table 1. All these tapes are tape-casted using plasticized PVB (polyvinyl butyral) as binder. A typical liquid adhesive for these tapes is 40% ethanol, 40% PPG, and 20wt% toluene solution, although details vary with each system. To produce solid coatings of a themally activated adhesive, a water-based suspension of ethylenelacrylic acid with 25% solids, is used. The average molecular weight and softening point of polymer are 8000 Amu and 42OC, respectively.

65) Ikbound at 550°C and fire at hi her temperature for densi f" ication.Fig. 1 Schematic of lamination of ceramic green tape using water-based emulsion of tough co-polymer that becomes soluble in PVB at 1 OO°CA dead load lamination procedure was used to study the lamination efficiency for non- liquid adhesive (coated tape), as shown in Fig. 1. After laser-cutting and stacking, the laminates were put into an oven at a low temperature (-125OC) under very low pressure (-2.5 kPa). Under these conditions, the solid polymer film softens, becomes tacky and bonds all layers together. Subsequently, the laminates were heated without any pressure to a higher temperature (175OC) that is still below the binder decomposition onset and held for longer time (5-10 hrs). The viscosity of polymer film decreases, and it diffuses into green tape at this point. Therefore, the interface between green tapes disappears. After firing, monolithic ceramic parts are obtained.

The viscosity of polymer film decreases, and it diffuses into green tape at this point. Therefore, the interface between green tapes disappears. After firing, monolithic ceramic parts are obtained. In the CAM-LEM machine, the coated film was heated to -1 50°C in a few seconds to melt the polymer film using radiantheat, then next layer was stacked and rolled to apply pressure. All the steps are robotically. The following pressureless annealling, debinding and firing are the same as shown in Fig. 1.Mass transportation during adhesive laminationIn prior work, a liquid-adhesive joining process was described that has proven very successful in terms of laminating a wide variety of ceramic green tapes. A variety of adhesive formulations have been developed for different feedstock. Solvents can be highly effective at assisting the lamination of ceramic green tape, but highly volatile solvent mixtures are difficult to control.

Solvents can be highly effective at assisting the lamination of ceramic green tape, but highly volatile solvent mixtures are difficult to control. For liquid adhesive lamination, the mass transportation during lamination is shown in Fig.2.In the experiments, it is found that the PPG component in liquid adhesive is very important for good lamination. Therefore, the interaction of PPG with ceramic green tape was studied [7]. Fig. 3 shows the DMA (dynamic mechanical analysis) test results of PPG-coated Coors' alumina green tape. For as-received tape, there are two peaks, corresponding two organic phases. The peak at 32OC corresponds to the T, temperature of binder and the peak at -7 1°C corresponds to the T, temperature of the plasticizer. After heating a coated green tape to 80°C for a prolonged time, the binder's T, shifted to lower temperature and the peak becomes broader.

This result indicated that the diluent, PPG, difhsed into binder.Evaporation of solventsDiffusion of solventsDeposited polymer filmCeramic green tapeSurface binder6a- layer thickness softened by adhesive6t- total thickness of green tapeFig.2 Mass transportation during liquid-adhesive laminationLoss Tangent 10"150.0100.050.0 Temperature (OC) 0.050.0100.0Fig.3 Dynamic mechanical behavior of Coors' alumina tapeCoated feedstock from commercial ceramic green tapeThe ability of the PPG to migrate suggested that a polymer film with a low melting (or softening) point may be used as adhesive for ceramic green tape lamination at elevatedTable 2 Lamination of polymer-coated Coors; alumina tapeLayers Thickness Weight12 7.15 10.83Lamination Temperature (OC) Time (hr) Pressure (kPa) Strain (%) Weight loss (%)80 2 2.4 -0.2 0.08Anneal at 175OC for 1 Ohrs Strain (%) Weight loss (%)0.70 3.89Debind and Bisque Fire at 800°C Strain(%) -0.98 Weight loss (%) 9.72Lamination Result (Fired at 1560°C, 0.5hr)delam2.26 9.72 good2.29 9.66 good2.44 9.69 good2.28 9.72 goodtemperature.

This was the underlying idea for the development at the solid system. The polymer coating on ceramic green tape was made by using a Doctor blade, as would be done in tape casting. Coated feedstocks were laminated using the procedure shown in Fig. 1. Preliminary results have already presented [5]. The detailed results are shown in Table 2. Most laminates are 12 layers. Different lamination temperatures, from 80°C to 175"C, were tested while keeping lamination time and pressure constant. It can be seen that, with increasing lamination temperature, lamination strain increased due to the lowering of viscosity of the polymer film and diffusion into the green tape. After pressureless annealing at 175"C, the strain of all of specimens except for that laminated at 80°C increased to 2.1 to 2.4%, which corresponds to the original polymer film thickness.

After pressureless annealing at 175"C, the strain of all of specimens except for that laminated at 80°C increased to 2.1 to 2.4%, which corresponds to the original polymer film thickness. At this time, it is evident that the polymer coating is diffused into green tape, because the strain after bisque firing at 800°C did not change much. According to the laminate thickness and strain, the polymer coating thickness can be calculated in the range of 14 to 16 pm. It can also be seen that the lamination weight loss is very small if lamination temperature is lower than 175°C. The total weight loss after pressureless annealing is 3.6 to 3.9%. At that time, most plasticizer in the binder, which has much lower molecular weight (300-600) than binder and coating polymer, was evaporated. The total organic component (binder and plasticizer) in ceramic green tape is 8.5wt%. The specimen laminated at 80°C has negative strain due to the buckling delamination.

The total organic component (binder and plasticizer) in ceramic green tape is 8.5wt%. The specimen laminated at 80°C has negative strain due to the buckling delamination. Therefore, lamination temperature should be 100°C or higher. However, the green tape will too soft if temperature is too high.Fig.4 shows a laminate of coated Coors' alumina tape. The cross-section of fired specimen subjected to a dye-penetrant indicates 100% lamination efficiency. No delamination was observed. Three layer atomizer components also were fabricated using coated Coors' alumina tape, shown in Fig.5.

No delamination was observed. Three layer atomizer components also were fabricated using coated Coors' alumina tape, shown in Fig.5. Fugitive tape, made from graphite and corn starch, was used in hollow area to ensure uniform pressure distribution during dead load lamination.(a)Edge of as-laminated specimen at 125°C for 2 hrs(b) Cross-section of fired alumina with penetrated ink (anneal at 175°C for 10 hrs, debound at 550°C, and fired at 1 560°C for 30 minFig.4 13-layer laminate of coated Coors' alumina (one inch square in green state)Coated feedstocks of fine ceramic green tapesCoated feedstock were also prepared from in-house alumina and silicon nitride ceramic green tapes.

The two kinds of feedstock were laminated using polymer films of 18 and 28 pm, respectively, at 125°C and then pressureless annealed at 175OC for 15 hrs.( a ) As-la~i~inatedat 125°CFired at 1560°C fbr 30 n ~ i nFig.5 3-layer atomizer made from coated Coors' alumina tapetapc at 1 25°C for 2 hrs( c ) As-laminated coatcd-silicon nitride tapc at 1 2S°C for 2 hrs( d ) Anneal ( c ) at 175°C for 15 hrsFig.6 Cross-section of laminates of coated fine ceramic green tapesThe cross-section of laminates are shown in Fig.6. It can be seen that, unlike Coors' alumina tape, a thin polymer film (3-4 pm) still exists between green tape after 15 hrs annealing at 175OC. The reason is that the polymer coating is too thick. With the diffusion of polymer coating into green tape, the pores at near surface of green tape become saturated and a dense layer was formed at near surface, which is apparent as discoloration.

With the diffusion of polymer coating into green tape, the pores at near surface of green tape become saturated and a dense layer was formed at near surface, which is apparent as discoloration. Therefore, continued diffusion of the polymer coating into the green tape will meet high resistance and the diffusion rate will be very slow. It may take longer time (20 or 30 hrs) for polymer coating to diffuse into green tape completely, which is necessary to avoid delamination during firing.(b) t(pn1 coating( c ) 3 prn coatingFig.7 Thinner polymer coating on in-house alumina tape(a), (b) As-lamirlatcd at 125°C for 2 hrs(c), (d) Heated to 175OC and hcld for 1 0 hrsFig.8 10-layer laminate of coated half-micron alumina tapeTherefore, thinner polymer coatings were prepared by diluting solid concentration in polymer emulsion. Coating thickness can be controlled in the range of 3 to 8 pm, as shown in Fig.7.

Coating thickness can be controlled in the range of 3 to 8 pm, as shown in Fig.7. Coated half-micron alumina tape was laminated at 125OC and then pressurelessly annealed at 175OC. SEM images of cross-section of the laminates are shown in Fig.8. It can be seen that, after annealing at 17 5*C, polymer coating diffused into green tape completely, and the interface between layers disappeared. Alumina atomizers were also fabricated by CAM-LEM machine. All the lamination steps were then finished robotically. Good lamination was obtained.ConclusionsLiquid-adhesive lamination is a mature technology and has been successfully used to produce advanced ceramic parts. Non-liquid adhesives have been developed and demonstrated. Though solid and nonsticky at room temperature, at low temperatures, it softens and bonds layers together.

Non-liquid adhesives have been developed and demonstrated. Though solid and nonsticky at room temperature, at low temperatures, it softens and bonds layers together. As temperature is increased, the adhesive becomes fluid and is dissolved in the binder.(3) Fine particle size green-tape require very controlled thin coating.References1. P. Wei, J.D. Cawley, and A.H. Heuer, CAM-LEM Processing of Ceramic Components - Lamination Technology, Proc. 6th Eur. Conf. Rapid Prototyping Manuf., edited by P.M. Dickens, Uni. of Nottingham, 1997,95.2. J.D. Cawley and Zhien Liu, Applying Tape Casting to Layered Manufacturing Processes, Ceramic Industry, 128[3]42(1998).3. Zhien Liu, P.Wei, A.H. Heuer; and J.D. Cawley, Metal and Ceramic Components Made via CAM-LEM Technology, Proc. Solid Freeform Fab. Symp., University of Texas at Austin, 1996, pp. 377.4. Zhien Liu; T-C KO; J. Best; J.D. Cawley, and A.H. Heuer, CAM-LEM Processing: Materials Flexibility, Proc. Solid Freeform Fab.

Symp., University of Texas at Austin, 1996, pp. 377.4. Zhien Liu; T-C KO; J. Best; J.D. Cawley, and A.H. Heuer, CAM-LEM Processing: Materials Flexibility, Proc. Solid Freeform Fab. Symp., University of Texas at Austin, 1997, pp. 379.5. J.D. Cawley; Zhien Liu; J. Mou, and A.H. Heuer, Materials Issues in Laminated Object Manufacturing of Powder-Based Systems, Proc. Solid Freeform Fab. S p p . , University of Texas at Austin, 1998, pp. 503.6. J.D. Cawley, B.D. Kernan, and W.H. Glime, US Patent Pending. 7. N.Suppakara, PhD Dissertation, Case Western Reserve University, 1999.Acknowledgements: Brian B. Mathewson at CAM-LEM, Inc. conducted the lamination trials with the CAM-LEM machine and produced the test lot at atomizer. Funding for yhis work was provided through a NASNCWRU Cooperative Agreement on Ceramic Processing (NCC3-404)..

lour Rapid Prototyping based on SLS processLing Wai Ming and Ian GibsonCentre for Advanced Product Development Technologies Department of Mechanical Engineering The University of Hong Kong Pokfulam Road, Hong KongAbstract Currently, the colour of parts made by the Selective Laser Sintering (SLS) process depends on the colour of the material used. SLS cannot make multiple coloured prototypes because only homogeneous powder (or powder composites) can be used during the process. In this paper, an ink-jet based mechanism is designed to print multiple colours onto the prototypes. The surface tension of conventional ink used in bubble jet printers is so high that it cannot penetrate into the SLS powders. To reduce the surface tension, ethanol is added to it. The ratio of ink to ethanol for necessary penetration is determined by experiments. This paper will go on to describe how the ink affects the structure of SLS materials.

The ratio of ink to ethanol for necessary penetration is determined by experiments. This paper will go on to describe how the ink affects the structure of SLS materials. The effect of temperature on the penetration of ink into powders will also be discussed.Keywords: Rapid Prototyping (RP), Selective Laser Sintering (SLS), Colour RP, Multiple material RP.Introduction Selective Laser Sintering (SLS) is a RP process for converting computerized 3D solid designs to physical objects by using a layered, powder-based laser manufacturing method [1]. Over the last 10 years, it has been shown that SLS can process a variety of materials with new materials showing a particular improvement in part accuracy and strength aimed at increasing SLS's range of functional applications. Future development of SLS appears to be focusing on rapid tooling and direct manufacturing. Many direct manufacturing applications will benefit from the incorporation of colour into the models.

Future development of SLS appears to be focusing on rapid tooling and direct manufacturing. Many direct manufacturing applications will benefit from the incorporation of colour into the models. In 1998, Michael Rees discussed the issue of colour RP and stated a number of reasons for its implementation [2]. Colour RP has already been used to highlight features in medical and other models. It is an additional tool for communication of concepts and ideas; the contrasting effect of different colours provides a means for distinguishing features in a much easier way than when using monochromatic processes. Currently, it is only possible to make coloured SLS prototypes using coloured powders or to use coloured infiltrants or coatings at the post-process stage.

Currently, it is only possible to make coloured SLS prototypes using coloured powders or to use coloured infiltrants or coatings at the post-process stage. With the aim of producing an automated, in-process method of creating multiple-coloured SLS prototypes, a colour printing system is described and discussed in this paper.Colouring using Ink-jet technology Ink-jet is a versatile, non-contact printing technique, which has grown dramatically in popularity over the last ten years. It has the ability to print fixed and variable data at high speeds to virtually any surface, porous and non-porous, including those that are irregular and fragile [3]. A wide range of colours can be produced by this technology. It satisfies the basic conditions for colouring SLS prototypes since all SLS parts and materials exhibit porosity.227Figure 1 shows an ink jet printing mechanism designed to fit inside the Sinterstation 2000 machine.

The main body of the design is the printing mechanism taken from the Canon 2000 bubble jet printer. This is incorporated into a 2-dimensional plotter type system, which has a swing arm on which the printing mechanism is mounted. During printing, this swing arm will move to the top of the part bed and jet ink onto the powder selectively. It will then move out of the way of the roller mechanism that is used to spread powder onto the part bed.Figure 1. Ink-jet printing mechanism fits inside the Sinterstation 2000 machineExperiments have been performed inside the Sinterstation using the Trueform material. The printing mechanism is not designed to run at the high temperature required for processing the Duraform material.

The printing mechanism is not designed to run at the high temperature required for processing the Duraform material. The standard operating temperature of Trueform is 68(cid:176) C, and therefore can be used for experiments.Wetting of ink The print quality on the surface of the powder will depend on the dispersion of the ink, which in turn relates to the wetting of a solid by a liquid. Wetting is a surface phenomenon [4 & 5]. When a drop of liquid falls on to a solid surface, wetting can be determined by the contact angle between the droplet and the solid surface (figure 2).Figure 2.

When a drop of liquid falls on to a solid surface, wetting can be determined by the contact angle between the droplet and the solid surface (figure 2).Figure 2. Angle of contact with the solid surfaceAssuming the various forces can be represented by the surface tensions acting in the direction of the surfaces, then/ AS=/ LS+/ ALcos(1)g S/A, g S/L and g L/A are the interfacial tensions of the solid-vapour, solid-liquid, and liquid-vapour interfaces respectively.228Combining this with the Dupré equation [6]= LSW /+ / AS/ AL/ LS(2)Where WS/L is the work of adhesion of a solid for a liquid.Equation (1) becomesLSW /=( 1/ AL+cos)(3)Thus if the forces of attraction between liquid and solid are equal to or greater than those between liquid and liquid, the angle of contact will be close to zero. If the adhesion of liquid to the solid is less than that to the liquid itself, the contact angle will be finite.

If the adhesion of liquid to the solid is less than that to the liquid itself, the contact angle will be finite. The wetting of solid is said to be perfect, which means that the liquid tends to disperse and penetrate, when the contact angle is zero. The solid is partially wetted when the contact angle is finite. If q is greater than 90(cid:176) , the solid is not wetted, and the liquid droplet is not easy to disperse. It is completely not wetted when the contact angle is 180(cid:176) . Therefore prediction of contact angle can give a quantitative idea of dispersion and penetrationPenetration of ink Ink penetration is another factor in controlling the print quality on SLS prototypes.

Deep penetration will mix the colours together, whilst shallow penetration may make a ‘gap’ between layers of powder.If we assume powder consists of small capillaries of constant radius R, then the penetration rate of a liquid into powder is given by the Lucas-Washburn Equation [7] below:2 L=ÆRcos 2t(4)L is the penetration distance after time t, R is the pore radius, g is the surface tension, q contact angle, t is the time and h is the liquid viscosityThis equation has been used in Three-dimensional Printing [8] to determine the infiltration of the binder into the powder.

There should be some corrections in this equation to account for the granular structure of the powder, but it is significant only for large penetration depths and it is neglected in this discussion.In the case where there is no interaction between the liquid and the powder, and the same volume of liquid is always added to the powder, the penetration depth of the liquid should be constant for all times.

As the pore radius R is a constant parameter, by rearranging equation 4:t=2 RÆ2 Lcos(5)229is thet= KÆcos(6)As K = 2•L2/R = constant, then= KtVSC(7)VSC is the ratio of h/(g• cosq )The direct relationship between the penetration time and VSC (for Viscosity, Surface tension and Contact angle) ratio is a good indication of liquid dispersion ability.Ink penetration experiment A preliminary investigation of ink penetration on powder involved inks of primary colours (cyan, yellow and magenta) and two powders, Trueform and Duraform Polyamide (simply called Duraform). The inks are produced by the Acujet Company (Los Angeles, California), and numbered AJ-211, AJ-212 and AJ-213 respectively. The specific gravity and powder density of Trueform and Duraform Polyamide powders are [9 & 10]:Duraform TrueformSpecific Gravity, 20(cid:176) C (g/cm3) 0.97 0.95Powder Density, Tap (g/cm3) 0.59 0.35Table 1.

Specific gravity and powder density of Trueform and Duraform powdersSih [11] defines the porosity of powder e as:=solidbulk(8)solidwhere rsolid is the density of the powder particle and r bulk is the bulk density of the powderFrom this equation, we find that the porosity of Trueform and Duraform are 0.6316 and 0.3918 respectively. Assume that there is no interaction between the ink and the powder, and there are no pores in the powder particle. The only way that the ink can penetrate is through the pores between the powder particles. Liquid should penetrate faster into the powder of higher porosity. Therefore the penetration time of ink in Trueform should be faster than that in Duraform.Contact angle of the ink drop on the powder surface was measured by direct measurement method [4].

Therefore the penetration time of ink in Trueform should be faster than that in Duraform.Contact angle of the ink drop on the powder surface was measured by direct measurement method [4]. Figure 3 shows a schematic of a micro-lens (Pulnix TM-765) which is connected to a digital camera that can give up to 25 times magnification of the object and therefore allows a direct evaluation of the contact angle.230Figure 3. Apparatus used for a direct measurement of contact angleThe viscosity and surface tension of the ink were measured using viscometer and surface tensiometer respectively.The penetration rate of ink into the powder was measured by recording the time required for a 0.05 ml drop to enter a loosely packed bed of powder. By using a micropipette, the drop was ejected to the surface of the powder.

By using a micropipette, the drop was ejected to the surface of the powder. It was assumed the drop penetrated fully if the liquid meniscus could no longer be observed.Initial tests found that primary inks (cyan, yellow and magenta) would not penetrate to either the Duraform or Trueform powders. Figure 4 shows the magenta ink drop on the surface of Duraform powder. It is clearly seen that the contact angle of the drop is greater than 90(cid:176) . It is not easy for the ink to disperse and cannot penetrate the powder.Figure 4. A magenta ink drop on the surface of Duraform powderTo remedy this problem the ink was mixed with ethanol (UN No. 1170) in the ratios 100 to 10, 100 to 20, 100 to 40, 100 to 60 and 100 to 80. In experiments using magenta ink on Trueform and Duraform respectively, the penetration rate increases with the ink/ethanol ratio (Figures 5 & 6).

In experiments using magenta ink on Trueform and Duraform respectively, the penetration rate increases with the ink/ethanol ratio (Figures 5 & 6). This agrees with the relationships between the ink/ethanol ratio and the VSC ratio derived from equation (7) (Figures 7 & 8). When cyan ink and yellow ink were used instead, similar results of penetration rate are obtained. The different penetration rates between primary inks are due to their different original viscosity, contact angle and surface tension values.Figure 5 & 6. Penetration time against ink/ethanol ratio (Figure 5: Duraform. Figure 6: Trueform)231Figure 7 & 8. VSC ratio against ink/ethanol ratio (Figure 7: Duraform. Figure 8: Trueform)Discussion A colour Three Dimensional Printing system from Z-Corp is expected to come onto the market at the end of the year 2000. The principle of the mechanism is to use an ink-jet based printer to add colour ink on to the powder selectively.

The principle of the mechanism is to use an ink-jet based printer to add colour ink on to the powder selectively. The ink also acts as an adhesive to bind the powder together. The printing mechanism discussed here uses the same principle but without the need to bind the particles together. The SLS process also has the advantage of being able to produce prototypes that have sufficient strength for functional applications like direct manufacturing.Ethanol is an additive used to make pure ink penetrable to the powder. It will also help the ink to disperse. The amount of ethanol in the ink should be carefully chosen. If the amount is insufficient, the ink will take too long to penetrate into the powder. Inversely too much will cause destruction of the powder structure. Figure 9 shows an SEM photo of Trueform powder particles after the addition of pure magenta ink. It is seen that some of the particles are bounded together.

Figure 9 shows an SEM photo of Trueform powder particles after the addition of pure magenta ink. It is seen that some of the particles are bounded together. Pure ink mixed with a small amount of alcohol will begin to dissolve the Trueform particles. The more ethanol used in the ink, the more the Trueform particles will be dissolved. Similar results have been observed when Duraform powder is used, with the additional effect of the Duraform particles being reshaped (Figure 10).Figure 9. Coloured Trueform particles. Figure 10. Coloured Duraform particles.The penetration rate of the inks into the Trueform and Duraform powders is consistent with the Lucas-Washburn equation. However, this equation cannot predict the actual penetration time. It is correct only when there is no interaction between the solid and the liquid, and it has been shown that there is dissolution of powder by the ink.

It is correct only when there is no interaction between the solid and the liquid, and it has been shown that there is dissolution of powder by the ink. Despite this it is a good equation to indicate the trend of ink penetration.232Experiments were carried out at the temperature of 20(cid:176) C. The Sinterstation 2000 machine generally runs at higher temperatures. Even the low temperature Trueform has the part bed set to 68(cid:176) C. Duraform has the part bed set to 182(cid:176) C. With such high temperatures, there should be some changes in the viscosity and surface tension of the ink. Penetration experiments were conducted at different temperatures and the results reveal that there is a similar trend between penetration time and ink/ethanol ratio for different temperatures (Figures 11&12).Figure 11 & 12.

Penetration of ink against ink/ethanol ratio at the sintering temperature of SLS material (Fig 11: Duraform Fig 12: Trueform)Adding ink appears to improve the surface of prototype. Figure 13 shows surface roughness profiles on a coloured Trueform surface. Ra is the arithmetic mean of the departures of the profile from the mean line, and Rt is the maximum peak to valley height of the profile in the assessment length. When compared with figure 14, the surface roughness profiles on Trueform surface without colour, it is shown that both Ra and Rt are greatly decreased after adding colour.Figure 13. Surface roughness profile on coloured Trueform surface (Ra = 4.9397m m, Rt = 34.2077m m, horizontal scale 500m m/division)Figure 14. Surface roughness profile on uncoloured Trueform surface (Ra = 13.4728m m, Rt = 98.6471m m, horizontal scale 500m m/division)The diameter of the ink droplet used in all the above experiments is about 4-6mm.

It is much larger than the size of the pores (ranges from 30m m to 60m m). If the diameter of the ink droplet is smaller than the pore size, ink droplets can go directly to the pores233and the penetration time should become faster. The diameter of the ink jetted from Canon BJC 210 bubble jet printer is about 100m m, which is larger than the pore size. Therefore the result from the penetration experiments should be the same in the real ink-jet printing situation.Conclusion It has been shown that a colour SLS process is feasible using the application of ink-jet printing. It is necessary to determine a range of ink/ethanol ratio that can be used in colouring SLS prototypes. In this range, the ink can penetrate to the powder and disperse away at an acceptable time, and it should affect the structure of the powder to the minimum extent. This will be studied by further experiment.References 1.

The "Tailored Forming” process chain consists of a deposition welding process, hot forming, machining and an optional heat treatment. This paper focuses on the combination of laser hot-wire cladding with subsequent hot forming to produce hybrid components. The applicability is investigated for different material combinations and component geometries, e.g. a shaft with a bearing seat or a bevel gear. Austenitic stainless steel AISI 316L and martensitic valve steel AISI HNV3 are used as cladding materials, mild steel AISI 1022M and case hardening steel AISI 5120 are used as base materials. The resulting component properties after laser hot-wire cladding and hot forming such as hardness, microstructure and residual stress state are presented. In the cladding and the heat-affected zone, the hot forming process causes a transformation from a welding microstructure to a fine-grained forming microstructure.

In the cladding and the heat-affected zone, the hot forming process causes a transformation from a welding microstructure to a fine-grained forming microstructure. Hot forming significantly affects the residual stress state in the cladding the resulting residual stress state depends on the material combination.Keywords: Laser hot-wire cladding, Cladding, Hot forming, Residual stress, Microstructure, Hardness, Tailored FormingIntroductionMetallic claddings are used in a wide range of applications to protect the component from environmental influences or high loads. The cladding material is adapted to the type of load applied to the hybrid component. Claddings can protect the base material from chemical, physical and/or tribological stresses. For example, a cladding of stainless steel can protect the base material from corrosion, while materials with high hardness can protect the base material from abrasive wear [1].

For example, a cladding of stainless steel can protect the base material from corrosion, while materials with high hardness can protect the base material from abrasive wear [1]. Due to the hybrid design with load-adapted claddings, the component can be made from low-cost mild steel and the higher-priced, high-strength or corrosion-resistant alloys are only used selectively in stressed component areas. Accordingly, material costs can be reduced by such resource-efficient material selection.State of the artA number of processes are available for the application of claddings, which differ in the properties produced. For the manufacturing of claddings with a thickness from 0.5 mm to several millimeters, cladding processes based on welding are used. In most cases, the application of a single weld seam is not sufficient to achieve a wide cladding. By applying overlapping weld seams, a sufficiently wide cladding is produced [2].

In most cases, the application of a single weld seam is not sufficient to achieve a wide cladding. By applying overlapping weld seams, a sufficiently wide cladding is produced [2]. The surface of the cladding has a high waviness and roughness, which is why machining is usually used after cladding to achieve the application-specific surface quality [3]. In order to keep the machining effort as low as possible, a number of investigations were carried out and models were developed which predict the waviness perpendicular to the welding direction depending on process parameters of the deposition welding process and can thus be used to optimize the rework effort [4–6].1223The microstructure and hardness of additively manufactured parts can be adapted to the application by subsequent heat treatment. Junker et al.

Junker et al. investigated the influence of heat treatment strategies on additively manufactured parts of hot work tool steel and compared the results to conventionally manufactured samples [7]. The hardness and compression strength after deposition welding was similar to those of conventionally manufactured samples after hardening but the yield strength was lower. After heat treatment, consisting of austenitization, quenching and three tempering cycles, the mechanical properties of the additively manufactured samples were similar to those of the conventionally manufactured sample. Tempering without pre-austenitization of the additively manufactured sample led to an improvement of the mechanical properties exceeding those of conventionally manufactured samples.Hafenecker et al. investigated the cold forming of additively manufactured cylinders of AISI 316L and of Ti-6Al-4V by a cup backwards-extrusion process [8].

investigated the cold forming of additively manufactured cylinders of AISI 316L and of Ti-6Al-4V by a cup backwards-extrusion process [8]. The cylinders were manufactured by a laser-based powder bed fusion process. The surface roughness of the lateral surface of AISI 316L samples was reduced by 90 % by cold forming and geometric shapes like sharp edges were improved. Due to the low formability, no improvement could be seen for the Ti-6Al-4V samples.In laser deposition welding, residual stresses occur due to the high temperature gradients and the rapid cooling of the melt pool, which can lead to distortion or crack formation depending on the component geometry and material combination. High tensile residual stresses usually occur near the surface, which can have a negative effect on the service life of the component. Various strategies can be used to reduce residual stresses or to selectively set beneficial residual stresses.

Various strategies can be used to reduce residual stresses or to selectively set beneficial residual stresses. Beneficial residual stresses are those that counteract load-induced stresses in the component. For example, by inducing compressive residual stresses into the surface and subsurface area, the service life of rolling element bearing raceways can be improved. The tensile residual stresses can be shifted towards the compressive range by processes that apply compressive mechanical loads like laser shock peening, deep rolling, hard turning or grinding. In the case of the latter two processes, it should be noted that additional tensile residual stresses can be induced by thermal effects. Residual stresses can be reduced by preheating the substrate or by subsequent heat treatment [9].In addition to the aforementioned processes of cold forming, as well as hardening and machining of cladded components to improve their properties, hot forming can also be used.

For hot forming, components are heated above the material specific recrystallisation temperature. After forming, the component is cooled, which can be done in ambient air or liquids, such as oil or water. Microstructure, residual stress state and hardness of the component are influenced during hot forming. Compared to cold forming, higher degrees of forming can be achieved, by which a higher change in shape of the component is reached [10].For the assessment of the suitability of the hot forming process for the further processing of cladded components, knowledge of the influence of the hot forming process on the surface and subsurface properties is required.

Hardness, residual stress state and microstructure of the cladding material have a significant influence on the application behaviour of hybrid components and will therefore be examined in detail in this study.Aims and perspectiveA process chain for the manufacturing of hybrid components consisting of a joining or cladding process, a hot forming process, a machining process and, if necessary, heat treatment is referred to as a Tailored Forming process chain. In this study, the focus is on the first two process steps of the Tailored Forming process chain, cladding and hot forming. In the following, the findings on resulting surface and subsurface properties after hot forming of semi-finished products cladded by means of laser hot-wire cladding are presented using the example of two geometries, a shaft with bearing seat and a bevel gear, and different material combinations. Hence, the following scientific questions have to be answered:1.

Hence, the following scientific questions have to be answered:1. How does the hot forming process affect the surface and subsurface properties of the hybrid component regardinga. Hardness b. Microstructure c. Residual stress state d. Cladding layer distribution?12242. How does hot forming affect the amount of machining required to achieve the final geometry of the hybrid part?Hybrid components of the aforementioned geometries are manufactured for the investigation of the researchquestions. Various material combinations are used for this purpose.Materials and MethodsTwo different component geometries are considered for the investigation. A cylindrical shaft is used as the base material for the production of hybrid shafts with bearing seats. The cladding is applied in the area of the bearing seat. A shaft with a conical shape is used for the manufacturing of hybrid bevel gears.

The cladding is applied in the area of the bearing seat. A shaft with a conical shape is used for the manufacturing of hybrid bevel gears. The cylindrical shafts are made of mild steel AISI 1022M or case hardening steel AISI 5120 while the shafts with conical shape are made of AISI 1022M. Figure 1 shows an overview of the process steps that were investigated, as well as the component geometries and material combinations. The base material is shown in Figure 1 (A1) and (B1), the component geometry after cladding is shown in Figure 1 (A2) and (B2) and the hybrid component after hot forming is shown in Figure 1 (A3) and (B3).

The process steps and material selection are explained in detail below.Figure 1: Overview of the investigated process steps, material combinations and geometries: (A1) base material for bevel gear, (B1) base material for shaft with bearing seat, (A2) hybrid semi-finished part for bevel gear, (B2) hybrid semi-finished part for shaft with bearing seat, (A3) hybrid bevel gear after hot forming, (B3) hybrid shaft with bearing seat after hot formingThe hybrid semi-finished parts are manufactured by laser hot-wire cladding (LHWC). For this purpose, the MK-II coaxial deposition welding head, manufactured by Laser Zentrum Hannover e.V., is used. In this welding head, the laser beam is split into four partial beams and deflected outwards. The partial beams are reflected by mirrors and converge again into one common spot. A hot current source is used to preheat the welding wire. A diode laser is used as the laser beam source.

The partial beams are reflected by mirrors and converge again into one common spot. A hot current source is used to preheat the welding wire. A diode laser is used as the laser beam source. The base material is clamped in a rotary axis for the cladding process. By superimposing rotational movement and linear movement of the shaft and the processing head, spiral weld seams are applied to the shafts. The experimental setup is shown in Figure 2. Two different cladding materials are investigated. The AISI 316L austenitic stainless steel and the AISI HNV3 martensitic valve steel are used for the experiments. The AISI 316L stainless steel is characterized by excellent corrosion resistance, but has comparatively low strength. The material AISI HNV3 is characterized by high hardness and is used particularly in workpiece areas subject to high mechanical loads. The cladding materials investigated have different coefficients of thermal expansion.

The cladding materials investigated have different coefficients of thermal expansion. The mean coefficient of thermal expansion of AISI 316L between 20 and 400 °C is 17.5  10-6 K-1, which is significantly higher than the coefficient of thermal expansion of AISI 1022M (14.1 ⋅ 10-6 K-1) and AISI HNV3 (11.8 ⋅ 10-6 K-1).1225Figure 2: Experimental setup for laser hot-wire claddingIn order to study the effect of the forming process on the residual stress formation, shafts of the material combination AISI 1022M and AISI HNV3 are manufactured and compared with shafts of the material combination AISI 1022M and AISI 316L. By this, the influence of the thermal expansion coefficient on the residual stress formation can be determined. The influence of the cooling strategy on the hybrid components is investigated using shafts with bearing seats with the material combination AISI 5120 and AISI HNV3. After hot forming, the hybrid shafts are cooled in air or quenched in water.

After hot forming, the hybrid shafts are cooled in air or quenched in water. The high-strength AISI HNV3 cladding is used to study the bevel gear. The chemical composition of the cladding and base materials are shown in Table 1.Table 1: Chemical composition of the investigated materials [11–14]AISI 1022M AISI 5120 AISI HNV3 AISI 316LC 0.18-0.23 ≤0.4 0.17-0.22 ≤0.4 3.14 0.44 0.873 1.709 0.013SiMn 0.30-0.90 ≤0.03 ≤0.015 ≤0.30 1.10-1.40 0.4PSCr- 0.017 0.01≤0.035 1.00-1.20 -9.1 18.3Ni ≤0.30 - - 12.22Mo ≤0.08 - 0.01 2.641The cylindrical shafts are provided with a cladding of two layers and the conical shafts are provided with claddings of four layers. The geometry of the conical shaft is symmetrical, so that two claddings can be applied on one conical shaft. This design reduces the amount of waste when cutting the cladded conical shaft to size prior to die forging. Two bevel gears can be manufactured from one cone shaft in the subsequent forming process.

This design reduces the amount of waste when cutting the cladded conical shaft to size prior to die forging. Two bevel gears can be manufactured from one cone shaft in the subsequent forming process. An overview of the parameters used to generate claddings is given in Table 2. For some material combinations, the welding parameters were adjusted from the first to the subsequent layers due to the increase in temperature of the component. In these cases, both parameter values are given in the table.1226Fe Bal. Bal. Bal.

In these cases, both parameter values are given in the table.1226Fe Bal. Bal. Bal. Bal.Table 2: Parameters for LHWC of the hybrid parts with different material combinationsShaft with bearing seat AISI HNV3 AISI1022M 2.1/1.95 90/80 2 1,200 1.0 6 8ParameterUnitAISI 316L AISI 1022M 2.1 70 2AISI HNV3 AISI 5120 2.1 90/80 2 1,000 1.0 6.9 6Laser power PL Current (wire preheating) I Wire feed rate vw Welding speed vs Wire diameter dw Stickout lw Shielding gas flow 𝑉̇ (Argon) Laser beam source Focal length (collimation unit) fc mmkW A m/min mm/min 1,200 Mm mm l/min -1.0 6.5 6LDM 3000-40Bevel gear AISI HNV3 AISI 1022M 1.5/1.35 80 2 1,200 1.0 7.7 810063After cladding, the cylindrical shafts are formed into shafts with a bearing seat by means of cross-wedge rolling (CWR). CWR is a preforming process in which the masses are redistributed along the axis [15].

CWR is a preforming process in which the masses are redistributed along the axis [15]. Two wedge-shaped tools move in opposite directions and reshape the workpiece producing axially symmetrical geometries. The CWR module at the IPH – Institut für Integrierte Produktion gGmbH is a CWR machine with two flat tools. The velocity of the tools is set to 150 mm/s. The tools are preheated to 150 °C to reduce the loss of heat from the workpiece during forming. Prior to forming the workpiece is heated to 1,250 °C using an induction oven. The principle of CWR is shown in Figure 3. Hot forming is followed by cooling in air or quenching in water.Figure 3: Process of cross-wedge rolling of hybrid partsThe conical shafts are formed into a bevel gear by die forging. The single-stage forming process is carried out using a screw press (Lasco SPR 500, 40 kJ).

The single-stage forming process is carried out using a screw press (Lasco SPR 500, 40 kJ). To prevent tool failure due to the fluctuating semi-finished volumes caused by the weld seams, the forging tool has a compensating system of closure plate in combination with disk springs. This design allows processing of welded semi-finished products without prior machining of the surface. The bottom die was heated to a temperature of 200 °C with a heating collar. Both, the lower die and the punch, were manually coated with a graphite-based lubricant (AirForge 5308 from Oelheld). To form the bevel gear, the semi-finished products are heated for 20 minutes to a temperature of 1,200 °C in a furnace with circulating air and inserted into the die. Directly after forming the glowing bevel gear was transferred to a process- integrated heat treatment which is explained in [16].

Directly after forming the glowing bevel gear was transferred to a process- integrated heat treatment which is explained in [16]. The experimental setup is shown in Figure 4.1227Figure 4: Experimental setup for die forging of hybrid bevel gearsHardness, geometry and microstructure of the cladding are analyzed after LHWC and after hot forming in the cross-sections of the hybrid parts. Residual stress measurements are carried out by X-ray diffractometry using the sin²-method. For this purpose, a two-circle X-ray diffractometer by GE Inspection Technologies of the type Seifert XRD 3003TT is used. The measurements are made with Cr K radiation with an applied high voltage of UA = 30 kV and a current of IA = 35 mA. Residual stresses were determined in axial (II) and circumferential () directions.

Residual stresses were determined in axial (II) and circumferential () directions. For further details regarding the residual stress measurement, please refer to [17,18].Results and DiscussionThe cladding layer distribution after LHWC and after hot forming is shown in the cross-sections in Figure 5. CWR of the hybrid shafts reduces the thickness and increases the width of the cladding. The cladding covers the area of the bearing seat. As a result, the highly stressed area of the shaft is provided with a suitable material while the low-stressed component area is made of the low-cost material. The teeth of the bevel gear are formed by die forging. The core of the teeth consists of the base material with good ductility, and the surface area with the cladding material has high-strength properties.

The core of the teeth consists of the base material with good ductility, and the surface area with the cladding material has high-strength properties. The tooth flanks, which are subjected to high loads during operation, thus have the necessary hardness for the loads during operation.Figure 5: Cladding layer distribution after LHWC and after hot forming: (a) cross-section of a hybrid shaft after LHCW, (b) cross section of a hybrid shaft after CWR, (c) cross-section of a hybrid bevel gear after LHWC, (d) and (e) cross-sections of a hybrid bevel gear after die forging1228The example of the shaft with bearing seat can be used to demonstrate the advantages of hot forming with regard to the amount of machining required and the resource efficiency. Figure 6 shows a visualization of the final geometry of the shaft with bearing seat after machining (blue).

Figure 6 shows a visualization of the final geometry of the shaft with bearing seat after machining (blue). If this shaft is produced by cladding with subsequent hot forming by CWR, the shaft has an allowance of around 0.5 mm after CWR, which has to be removed by machining (orange). It may be necessary to shorten the shaft at the ends (see Figure 1 (B3)), but this can be avoided by optimizing the base cylinder geometry. For this investigation, a long base cylinder was chosen to guarantee a complete formation of the shaft geometry, since the focus of the investigation was on the properties of the hybrid components. In further investigations, the base cylinder length can be optimized to ensure minimal rework. If the shaft is produced only by cladding and subsequent machining, the allowance over the shaft is not constant and the machining effort is significantly increased (green). More material has to be removed and is wasted.

More material has to be removed and is wasted. The amount of waste and machining can be reduced by the hot forming step by manufacturing a near- net-shape part. Reducing the amount of post-processing is of particular interest for materials with poor machinability. The extent to which a reduction in machining effort is possible depends on the target geometry.Figure 6: Visualization of the machining volume to reach the final geometry (blue) after CWR (orange) and after LHWC (green)In the following, the influence of hot forming on microstructure, hardness and residual stresses is demonstrated using the shaft with bearing seat as an example. The hardness curves of the shaft with bearing seat for the different material combinations after LHWC and CWR are shown in Figure 7. For better comparability of the hardness curves, the transition from cladding to base material was chosen as the zero point of the x-axis for the plot.

For better comparability of the hardness curves, the transition from cladding to base material was chosen as the zero point of the x-axis for the plot. The high-strength claddings have a hardness of 670 to 750 HV0.1 after deposition welding, while the corrosion-resistant AISI 316L cladding has a hardness of only 200 HV0.1. The hardness of the unalloyed carbon steel AISI 1022M ranges from 150 to 200 HV0.1 and the case hardening steel AISI 5120 has a hardness of 200 to 250 HV0.1.Figure 7: Mean microhardness of the hybrid component with different material combinations after LHWC (left) and after CWR with different cooling strategies (right)1229After hot forming the hardness of the AISI 316L is only slightly increased to 200 to 250 HV0.1. The hardness of the unalloyed base material AISI 1022M is in the range of 150 to 170 HV0.1. The samples with a cladding of AISI HNV3 cooled in air show larger hardness variations than the quenched samples.

The samples with a cladding of AISI HNV3 cooled in air show larger hardness variations than the quenched samples. The hardness of the quenched specimens is in the range of 700 to 750 HV0.1 while the hardness of the aircooled specimens is in the range of 650 to 750 HV0.1. A significant increase in hardness is not achieved during quenching, since the cooling rates during air cooling are already sufficient for the formation of a martensitic microstructure. The large hardness variations are caused by scale formation and surface decarburization. The cooling strategy has a great impact on the hardness of the case hardening steel AISI 5120. The hardness of the quenched sample is more than 200 HV0.1 higher than the hardness of the aircooled sample. A hardness of 450 to 500 HV0.1 is achieved in the quenched sample.

The hardness of the quenched sample is more than 200 HV0.1 higher than the hardness of the aircooled sample. A hardness of 450 to 500 HV0.1 is achieved in the quenched sample. This indicates that hot forming with air cooling has only a slight effect on the hardness values compared to the cladded condition, while quenching in water can achieve strong increases in hardness depending on the material. For the subsequent application as a bearing seat, a hardness of the cladding of 690 HV +150 HV is targeted. The target hardness is achieved with the quenched samples. For the air-cooled samples, the hardness values fluctuate and are partly below the target value.

The target hardness is achieved with the quenched samples. For the air-cooled samples, the hardness values fluctuate and are partly below the target value. For these samples, subsequent heat treatment is necessary to reliably achieve the target hardness.Figure 8: Microstructure after LHWC (a,c,f,h) and CWR (b,d,e,g,i); microstructure of the base material AISI 1022M and AISI 5120 shown in (a,b,c,d,e) and microstructure of the cladding material AISI 316L and AISI HNV3 shown in in (f,g,h,i)The microstructure after LHWC and hot forming is shown in Figure 8. After LHWC, a Widmannstätten microstructure is present in the base material AISI 1022M (Figure 8 (a)). The base material AISI 5120 has a1230ferritic-pearlitic microstructure (Figure 8 (c)). The claddings of AISI HNV3 and AISI 316L have a martensitic and austenitic microstructure, respectively. By applying the second layer of the cladding, the first layer is subjected to heat treatment.

The claddings of AISI HNV3 and AISI 316L have a martensitic and austenitic microstructure, respectively. By applying the second layer of the cladding, the first layer is subjected to heat treatment. Anisostropic microstructure alignment is formed in the first layer (Figure 8 (f) and (h)). After forming, the Widmannstätten microstructure in the base material is transformed into a fine-grained ferritic-pearlitic microstructure ((Figure 8 (b)). Recrystallization by hot forming ensures that there is no longer any anisotropy in the microstructure of the cladding material ((Figure 8 (g) and (i)). The microstructure formation in the AISI HNV3 cladding is independent of the cooling method, but the cooling method influences the microstructure in the AISI 5120 base material.

The microstructure formation in the AISI HNV3 cladding is independent of the cooling method, but the cooling method influences the microstructure in the AISI 5120 base material. Cooling in ambient air leads to the formation of a ferritic-pearlitic microstructure, while quenching in water leads to the formation of a martensitic microstructure, which contributes to the aforementioned increase in hardness ((Figure 8 (d) and (e)).A comparison of residual stress in AISI HNV3 and AISI 316L claddings after CWR shows that hot forming influences the residual stress state. The resulting stress state in the cladding depends on the material combination. The residual stress depth profiles after LHWC and CWR are shown in Figure 9. After LHWC, tensile residual stresses occur in the claddings at the surface, changing to the compressive region with increasing distance from the surface.

After LHWC, tensile residual stresses occur in the claddings at the surface, changing to the compressive region with increasing distance from the surface. After reaching a compressive residual stress maximum, the residual stresses shift back into the tensile range (AISI 316L cladding) or to lower compressive residual stresses (AISI HNV3).Figure 9: Residual stress depth profile after LHWC and after CWRThe hot forming process has a great influence on the residual stress state. Near-surface residual stress values cannot be measured due to the formation of iron oxide. Residual stresses were measured in the axial and circumferential directions; measurement in the radial direction was not possible with the selected measurement method. The formation of residual stresses after CWR is influenced by three superimposed factors. The deformation of the shaft in radial direction during CWR induces compressive residual stresses into the shaft.

The deformation of the shaft in radial direction during CWR induces compressive residual stresses into the shaft. The different coefficients of thermal expansion of the cladding and the base material lead to strain incompatibilities during cooling of the hybrid component. This results in the formation of thermophysically induced residual stresses. Since the hybrid components were cooled in air, an inhomogeneous temperature distribution occurred, which can also influence the residual stress formation in the surface and the subsurface area. The residual stresses1231in the AISI 316L cladding with a high coefficient of thermal expansion are in the tensile range in both, axial and circumferential directions, after hot forming. The low thermal expansion coefficient of AISI 1022M hinders the shrinkage of the cladding during cooling, and the strain incompatibilities thus lead to the formation of tensile residual stresses in the cladding.

After hot forming of the AISI HNV3 cladding, the compressive residual stresses below the surface decrease. Up to a depth of approx. 1.22 mm, compressive residual stresses are present, after which compressive and tensile residual stresses occur.This shows that the residual stress state after hot forming depends on the selected material combination and that strain incompatibilities between base material and cladding material are the most important factor influencing the residual stress state. If the thermal expansion coefficient of the cladding is greater than that of the base material, tensile residual stresses are formed in the cladding; in the other case, compressive residual stresses are formed. The residual stress state in the cladding after hot forming can thus be influenced by smart material selection.

The residual stress state in the cladding after hot forming can thus be influenced by smart material selection. In particular, the option to achieve compressive residual stresses in the high-strength AISI HNV3 cladding is of special interest for component areas with high mechanical/tribological loads, since an increased service life of the cladding can be achieved.Conclusion and OutlookIn summary, the combination of a cladding process with subsequent hot forming affects the properties of thehybrid component as follows:Coarse-grained microstructure and anisotropy in the microstructure are transformed into homogeneous fine-grained forming microstructure by the hot forming processFor the production of complex geometries such as the bevel gear, simple semi-finished products can be used for the cladding process. The complex geometry is only created in the forming step.

The complex geometry is only created in the forming step. The simple semi- finished geometry offers potential for the application of cladding processes with high deposition rates and reduces programming effort and the number of necessary handling axes.Hot forming influences the residual stress state in the cladding and, depending on the material combination, advantageous compressive residual stresses can be achieved.Depending on the material used, it is possible to harden the hybrid component by quenching from the forging heat.The workpiece volume which has to be removed by machining in order to achieve the final geometry can be reduced by mass pre-distribution during hot forming.Die forging was used to produce a hybrid bevel gear in which the core of the tooth is made of ductile material and the tooth flank surface is made of high-strength material.

In this way, a load-adapted material distribution was realized.In the next step, the service life of the hybrid components has to be investigated. For this purpose, the bevel gear and the shaft with bearing seat must be machined and, in the case of the air-cooled samples, heat treated. In comparison with cladded components manufactured without hot forming, the positive influence of hot forming on the properties of the hybrid component can be quantified. The investigations can be extended to other material combinations. The combination of an AISI 1022M base material with an intermediate cladding layer of AISI 316L and an outer cladding layer of AISI HNV3 is of particular interest.

Combinatorial Parametric Engineering Model for Solid Freeform FabricationJ.C. Boudreaux NIST/Advanced Technology ProgramAbstract: Fabricated parts are often represented as compact connected smooth 3-manifolds with boundary, where the boundaries consist of compact smooth 2-manifolds. This class of mathematical structures includes topological spaces with enclosed voids and tunnels. Useful information about these structures are coded into level functions (Morse functions) which map points in the 3-manifold onto their height above a fixed plane. By definition, Morse functions are smooth functions, all of whose critical points are nondegenerate. This information is presented by the Reeb graph construction that develops a topologically informative skeleton of the manifold whose nodes are the critical points of the Morse function and whose edges are associated with the connected components between critical slices.

This approach accurately captures the SFF process: using a solid geometric model of the part, defining surface boundaries; selecting a part orientation; forming planar slices, decomposing the solid into a sequence of thin cross-sectional polyhedral layers; and then fabricating the part by producing the polyhedra by additive manufacturing. This note will define a qualitative and combinatorial parametric engineering model of the SFF part design process. The objects under study will be abstract simplicial complexes K with boundary ∂K. Systems of labeled 2-surfaces in K, called slices, will be associated with the cross-sectional polyhedral layers. The labeled slices are mapped into a family of digraph automata, which, unlike cellular automata, are defined not on regular lattices with simple connectivities (cells usually have either 4 or 8 cell neighborhoods) but on unrestricted digraphs whose connectivities are irregular and more complicated.1.

IntroductionEngineering designs are complicated data sets which contain such essential information about industrial products as CAD data , material specifications, dimensioning and tolerancing data, surface condition specifications, and much else besides. The complexity of designs mirrors the structural complexity of products. Products are properly modeled as networks of interacting components, each of which separated from the others by a (physical or conceptual) boundary, and all of which are interconnected by links such that outputs of one component may be passed as inputs to others. Products are also stratified structures, that is, a component at one level (a parent) may be refined into networks of components (the children) at the next-lower level.

Products are also stratified structures, that is, a component at one level (a parent) may be refined into networks of components (the children) at the next-lower level. In the stratification process, the links of the parent component are preserved: every input of the parent must be an input to at least one child and every output an output of at least one child.Stratification also leaves the parent’s boundary in tact, allowing it to become acontainer of the boundaries the children. Boundaries drawn in this manner impose a nest structure on the product: at the highest level is the product itself, followed by its immediate descendants, and so on, until all of the boundaries have been accounted for. In the product-as-network model, stratification boundaries mark off key components which can be then be identified by symbolic “tags” or labels.

In the product-as-network model, stratification boundaries mark off key components which can be then be identified by symbolic “tags” or labels. This symbolic structure permits stratified product networks to form a framework with respect to which the design features of components are defined. Features of one component may be either synthesized from those of its descendant components in a defined stratification or inherited by virtue of its links to other (in some cases, all other) components. The multiplicity of these relational ties goes a long way in explaining the observed conceptual diversity among feature563categories. For example, functional features are inherited from above by considering the constraints imposed by the component's children.

For example, functional features are inherited from above by considering the constraints imposed by the component's children. Specifically, the functional capabilities and operational behavior of components may be modeled by transfer functions which, given time-sequences of inputs (and possibly information about the current state of the component), produce time-sequences of outputs (and possibly a new state) after a specified time lag. Mating features are both synthesized from below and inherited from above because of the requirements imposed by assembly and joining processes. Form features are those through which all of the (often) contending requirements of the part are finally reconciled and by which the connectivity-induced constraints are resolved.Lurking in the background is the difficult notion of design completeness, whichmight be roughly defined as data sets which are sufficient to determine all of the essential properties of the components which instantiate them.

However, since no fixed meaning can be attributed to this notion, a few illustrative examples might suggest lines for future study. First, given a model of the “local” context of the component in the product network, a complete design for that component will need to be elaborated to the point that it possible to decide with confidence that its instances would satisfy all of the requirements and constraints active in that context. Second, a complete design and an inventory of available manufacturing capabilities and plant layouts should yield either manufacturing process plans and material flow routes or a determination that the component lies beyond the available productive capabilities. These two examples show why the notion of complete designs is as difficult as it is!In any event, constructing engineering designs is a difficult and costlyundertaking.

These two examples show why the notion of complete designs is as difficult as it is!In any event, constructing engineering designs is a difficult and costlyundertaking. It has often been proposed that the thing to do is to reuse designs by “instrumenting” them by defining a system of parameters for them. This has proven to be a difficult undertaking in its own right. Parametric engineering is based on the premise that an adequate inventory existing designs can be used to create active design templates and that by “plugging” in reasonable values for the embedded parameters, it is possible to automatically generate novel variant designs.In an earlier note [4], it has been argued that it is reasonable to model SFF-eligible parts as 3-manifolds with boundary. The boundary consists of a possibly disconnected set of compact smooth 2-manifolds (topological surfaces). For example, the 2-sphere is the single boundary of the closed 3-ball.

The boundary consists of a possibly disconnected set of compact smooth 2-manifolds (topological surfaces). For example, the 2-sphere is the single boundary of the closed 3-ball. In the general case, the class of structures includes more complicated objects that have enclosed voids and tunnels. But the fact that every compact 3-manifold M can be triangulated by an abstract simplicial complex K whose geometric realization in Euclidean 3-space, written |K|, is homeomorphic to M suggests that it is possible to construct a parametric engineering model directly on discrete combinatorial structures. From the parametric engineering point of view, geometric part definitions are massively overdetermined.Moreover, this note will suggest that a major part of the burden can be carriedby even simpler mathematical structures, called directed graphs, shortened to digraphs.

A digraph is defined as a pair (V,E), where V is an unempty set of vertices, E is a set of edges, together with a pair of functions start,stop:E →V such an edge h is interpreted as an arrow from the vertex start(h) to the vertex stop(h). This definition admits a very large class of structures. For example, digraphs may have many (even infinitely many) distinct parallel and antiparallel edges between the same two vertices. Digraphs may also have loop edges for which the start vertex is the same as the stop vertex.564Stratified product networks are complicated enough to illustrate the technical usefulness of digraphs. The first step is straightforward: the vertices of the product diagraph are symbolically “tagged” components; and the arrows consist of both structural links which connect parents with their children and general links between components. The second step handles repeated stratifications and is only a bit more complicated.

The second step handles repeated stratifications and is only a bit more complicated. Suppose that A is any arbitrary digraph whose vertices are distinct from those of V and that v is any vertex of V, then the desired stratification may be obtained by replacing v by A and adding one or more arrows to parallel those which originally connected v. In effect, the addition of A induces a three-fold partition of the edges of the new digraph: interior edges whose start and stop vertices are both in A, exterior edges whose start and stop vertices are both not in A, and boundary edges one of whose start or stop vertices is in A and other is not in A. This procedure shows that products may be represented by series of digraphs linked to one another by successive stratifications.The full power and potential of the parametric engineering model is stronglyapparent in the context of SFF.

This style of manufacturing is based on additive manufacturing processes which may resolved into the following steps: (1) create a solid geometric model of the part, defining surface boundaries and identifying interior (material) regions from the exterior (void) regions; (2) select a part orientation and then form planar slices, decomposing the solid into a sequence of thin cross-sectional polyhedral layers; and (3) fabricate the part by producing all of the polyhedra by any of several methods, including: stereolithography, selective laser sintering, laminated object manufacturing, fused deposition modeling, and three dimensional printing. In classical manufacturing work is applied to the surface of the part either through deformation process or by material removal.

In classical manufacturing work is applied to the surface of the part either through deformation process or by material removal. But the primary advantage of SFF is that it reduces the 3-space shape of compact objects to a sequence of 2-space slices (and possibly some mild “skinning” conditions).The necessary topological details are sketched in Sections 2 and 3 (see [11] fordetails). In Section 4 below, I will describe how the notion of a digraph can be applied to define a class of automata, called digraph automata, which I suggest will provide a natural computational framework for a parametric engineering model for SFF.2. Simplicial Complexes, Pseudomanifolds, and Edge PathsAn abstract simplex is a finite unempty set of elements, called vertices. If s is a simplex containing q+1 vertices, then s is a simplex of dimension q, that is, a q-simplex.

If s is a simplex containing q+1 vertices, then s is a simplex of dimension q, that is, a q-simplex. Simplexes have a clear geometric meaning: a 0-simplex is a set containing a single vertex representing a point (with the usual mild abuse of notation, points and vertices are identified); a 1-simplex is an unordered pair of vertices representing an edge; a 2-simplex is an unordered triple of vertices representing a triangle; a 3-simplex is an unordered quadruple of vertices representing a tetrahedron; and so on.Every unempty subset of a q-simplex s is a p-simplex, for some dimension p≤q,which is called a face (resp., a p-face) of s. For example, if s is a 3-simplex, then s contains four distinct 2-faces, six 1-faces, and four 0-faces. Every (proper) subset of s is called a (proper) face of s.565An abstract simplicial complex (henceforth, complex) K is a set of abstract simplexes such that if simplex s is an element of K then all faces of s are elements of K.

A complex is determined by the simplexes that it contains. Every complex is partially ordered by the inclusion relation between its faces. Thus, the pair (K,⊂ ) is a digraph whose vertices are the simplexes of K and whose arrows link each simplex with its proper faces. Those simplexes of K that are not proper faces of other simplexes are called maximal faces of K. The set of all maximal faces completely determines the complex, which may be generated by closing the set of maximal faces under the formation of subsets. For example, if {1,2} is a maximal face, then let [1,2] be the set of its unempty subsets, that is, {{1,2},{1},{2}}.If K is a complex, then its dimension (written dim K) is -1 if K is empty, and q if Kcontains at least one q-simplex but no (q+1)-simplex. Suppose K is a complex, then the q-dimensional skeleton (henceforth, q-skeleton) of K, written Kq , is the set of all of the p- simplexes of K such that p≤q. The q-skeleton of a complex is itself a complex.

The q-skeleton of a complex is itself a complex. In particular, the 1-skeleton of a complex consists of its vertices and edges.A simplicial mapping is a function from the vertices of one simplicial complex to to the vertices of another such that the images of the simplexes of the first are simplexes of the second. Two simplicial complexes are isomorphic if there is a bijective simplicial mapping from one to the other such that the inverse mapping is also simplicial. Given a Euclidean n-space of suitably high dimension, the convex hull every set of (k+1) pointwise independent points is a geometric k-simplex. Every finite abstract simplicial complex is isomorphic to a geometric simplicial complex, called its geometric realization.Since exotic complexes will have no role in this note, some additional restrictionson complexes will be imposed. For example, only those complexes that have finitely many elements will be considered.

For example, only those complexes that have finitely many elements will be considered. These complexes are not only of finite dimension but are also locally finite in the sense that each of their simplexes is a face of at most finitely many other component simplexes. Suppose that K is a finite simplicial complex, then:(R1) dim K ≤ 3. (R2) K is a pure complex if all maximal faces of K are of equal dimension, that is, K is a q-complex and every simplex s in K is a face of at least one q-simplex in K. (R3) K is a q-pseudomanifold (with boundary), that is, K is a pure q-complex and every (q-1)-face of K is the face of at most two distinct q-simplexes of K.The first two restrictions insure that the complexes under study have no isolated or dangling points, edges or triangles. A simple example shows that restriction (R3) can be violated even if both (R1) and (R2) are satisfied.

A simple example shows that restriction (R3) can be violated even if both (R1) and (R2) are satisfied. Suppose that K # is the 3-complex whose maximal faces are 3-simplexes [4,1,2,3], [5,1,2,3], and [6,1,2,3] whose vertices are distinct. Then K # is both a finite and a pure 3-complex, thereby satisfying the first two restrictions, but obviously fails to satisfy the third, since the 2-simplex {1,2,3} is a proper face of each of them. K # is mildly pathological in the sense that it cannot be realized as a polyhedron in 3-space, which is one of the pathologies which (R3) is designed to remedy.Let K be a 3-pseudomanifold and s be any arbitrarily chosen 2-simplex of K, then there are only two possible cases: either s is common face of two distinct 3-simplexes or s566is a face of exactly one 3-simplex in K. In the first case, the 2-simplex is said to be an interior face of K. In the second case, the 2-simplex and all of its unempty subsets are said to be boundary elements of K.

In the first case, the 2-simplex is said to be an interior face of K. In the second case, the 2-simplex and all of its unempty subsets are said to be boundary elements of K. Let ∂K be the set of all such boundary elements. Thus, the restrictions induce a partition on the 2-simplexes of 3-pseudomanifolds. The restrictions also imply that ∂K is a 2-complex, that is, that ∂K is a finite pure 2-complex, and that every 1-simplex (edge) in ∂K is a common face of precisely two 2-simplexes, that is, ∂K is a 2-pseudomanifold without boundary.Imagine a walk through a 3-pseudomanifold: start at a vertex, say v, then stepacross any of the finitely many edges containing that vertex, say {v,w}, to the companion vertex w, at which point either stop or repeat the process with w as the new starting point. Such a walk can be unambiguously encoded by a list of ordered pairs of vertices <v,w> representing the case in which one started at v, stepped across {v,w}, and then landed on w.

Such a walk can be unambiguously encoded by a list of ordered pairs of vertices <v,w> representing the case in which one started at v, stepped across {v,w}, and then landed on w. A list of ordered pairs generated in this manner is called an edge path. If K is a 3- pseudomanifold, then let (cid:31)(K) be the set of all edge paths through K. This list is contained in the 1-skeleton of K and has important algebraic properties. For example, every edge path has an inverse, that is, a path which starts from the original stop vertex and traverses the list of ordered pairs in the reverse order, stepping across <w,v> whenever the original stepped across <v,w>, and stops at the original start vertex. A composition operator can be defined for any pair of edge paths such that the stop vertex of the first is identical the start vertex of the second.

A composition operator can be defined for any pair of edge paths such that the stop vertex of the first is identical the start vertex of the second. The resulting algebra is called the edge path groupoid (see [11], 134-139).If there is at least one edge path between every pair of vertices of K, then K is saidto be (edge) connected. If a q-pseudomanifold is not connected, then it resolves into finitely many component q-pseudomanifolds which are connected. Henceforth, all pseudomanifolds will be assumed to be edge connected.Other types of paths, each having an associated definition of connectedness, may also be defined. For example, there are walks through 3-pseudomanifolds in which one starts an edge and then steps across any one of the 2-faces containing that edge to another edge of that 2-face.

For example, there are walks through 3-pseudomanifolds in which one starts an edge and then steps across any one of the 2-faces containing that edge to another edge of that 2-face. The associated path structure is a natural extension of the one defined for edge paths.More generally, let P be any one of these paths, then P may be represented as adigraph whose vertices are the (q-1)-faces encountered along the path and whose arrows are the q-faces across which one steps.3.

Slices, Patches, and Reeb GraphsGiven the importance of thin cross-sectional polyhedral layers in SFF-stylemanufacturing, the next step is to build a combinatorial mechanism to represent them:A slice of a 3-pseudomanifold K is a 2-pseudomanifold with boundary X such that (S1) X is a subcomplex of K and ∂X is a subset of ∂K; (S2) X decomposes K into upper and lower subcomplexes Ku(X) and Kl(X) whose union is K , whose intersection is X, and whose vertices are disjoint except for the shared vertices of X; and (S3) if v is any vertex of Ku(X) and w of Kl(X), every edge path in (cid:31)(K) from v to w contains at least one vertex in common with X.567If X is a proper subcomplex of ∂K consisting of one or more 2-simplexes and their faces, then X is a slice of K and in this case Ku(X)=X and Kl(X)=K. All sets X of this kind will be called boundary slices of K. Boundary slices may be composed of multiple components, each of which is a connected 2-complex with boundary.

All sets X of this kind will be called boundary slices of K. Boundary slices may be composed of multiple components, each of which is a connected 2-complex with boundary. These components will be called (boundary) patches. Thus, every boundary slice resolves into one or more patches. Since 3-pseudomanifolds are required to be finite, there must be minimal patches, specifically all boundary 2-pseudomanifolds consisting of a single 2-simplex and its faces. Let P be any unempty set of boundary slices of K closed under set intersection. Then every unempty element of P is itself a boundary slice and the pair (P,⊂ ) is a digraph whose arrows connect slices with subslices. The pair (P,⊂ ) will be called a boundary system for K.A sequence of slices (X1,...,Xm) is a cross-sectional slicing (cs-slicing) of K just incase for all 1≤i<j≤m the slices Xi and Xj are disjoint and Xj is a subset of Kl(Xj), or equivalently Xi is a subset of Ku(Xj).

Keep in mind that cs-slicing is defined combinatorially and does not require the 3-pseudomanifold to be embedded in any Euclidean space.From the perspective of Morse theory (see [2] [6],[10],[11]), the most interesting cs-slicings are those which (1) begin and end with minimal patches and which (2) mark the critical regions of the 3-pseudomanifold under study, that is, regions in which either the number of components or their topological properties change. An example of a change of the second kind is the first appearance of a new boundary patch (or void surrounded by boundary elements) in a patch whose earlier stages had none. Let cs- slicings of this kind be called Morse slicings.

Let cs- slicings of this kind be called Morse slicings. In general, 3-pseudomanifolds admit of many Morse slicings.Given a 3-complex K and a Morse slicing (X1,...,Xm), it is possible to build acombinatorial variant of a Reeb graph (compare [9]):Under the assumption that the slicing has already been reduced by eliminating consecutive pairs of homeomorphic slices, then (RG1) the vertices of the Reeb graph are components of slices; and (RG2) the arrows of the Reeb graph link each component of slice Xi to all of the components of slice Xi+1 which descend from it.As in the smooth case, Reeb graphs concisely encode important topological properties of 3-pseudomanifolds: the vertices encode the components of slices, and the arrows encode the 3-pseudomanifolds (like cylindrical solids) which link a component of slice Xi with all of the components of slice Xi+1 from which descend from it. The 3-pseudomanifolds paired with Reeb graph arrows may have voids and tunnels.

The 3-pseudomanifolds paired with Reeb graph arrows may have voids and tunnels. However, Reeb graphs are less successful in capturing properties affecting the geometric realizations of 3- pseudomanifolds. For example, there are many ways to embed an encoded Reeb graph into 3-space which are not geometrically equivalent (isotopic) to one another. This has to do with the mathematics of knots in 3-space (see [1]).4. Digraph Automata568In this section, I will propose a class of automata, called digraph automata, whichI can provide a plausible computational framework for a parametric engineering model for SFF.

In particular, I propose that this framework be based upon a functional model of computation in which lists are used to specify the application of a function (the first, or head, expression of the list) to its arguments (the rest of the list), and in which the computational processes are carried out by entities called evaluators (see [3]).Though evaluators may differ in their implementational details, they are all founded on an interpretive mode of operation. The behavior of the evaluator is described as a potentially infinite loop:(E1) read expressions from sources of input expressions, called input streams; (E2) evaluate the expressions in the context of the symbolic environment env (a list of symbol/value pairs), modifying env as required; and then (E3) write the expression resulting from the evaluation to an output stream, as required.The set of expressions includes three main classes.

First, there is a collection of atoms, which includes such familiar elements as characters, strings, integers, and floating-point numbers. When the evaluator is given any of these expressions to evaluate, it returns the expression itself. Second, symbols are expressions that have a more complicated role: when given a symbol, the evaluation procedure tries to match the symbol with the entries in the symbolic environment env. If a match is found, value corresponding to the symbol is returned; but if there is no match, then the interpreter returns a message to that effect. Third, when the evaluator is given a list to evaluate, it assumes that the expression is the application of a function, which must be the first component of the list, to the interpreted values of the arguments, which must be the remaining components of the list.

The value obtained is then returned as the value of the original list.The organization and control of the symbolic environment is perhaps the singlemost important task of the evaluation process. Each evaluator has access to an environment and can modify it in the course of operation. Moreover, during the evaluation of an expression a new local environment may be created and then destroyed after the evaluation is completed. The process is to create a list of symbol/value pairs and then append that list to the current environment; and when the appended material is no longer needed, remove it. Symbolic environments may be nested within each other and access rights for different evaluators can be precisely controlled (see [3], 44-46 for details). Thus, the value of a given expression is a context-sensitive function of the inputs and the currently active symbolic environment.It’s now a short step to digraph automata.

Thus, the value of a given expression is a context-sensitive function of the inputs and the currently active symbolic environment.It’s now a short step to digraph automata. The general approach to be taken is todevelop a computational theory based on an agent-oriented model of computation. Agents are computing elements that have persistent internal states, which are able to sense and act upon their external environment, and which, if suitably configured, are able to communicate with one another. Agents which share the same communication conventions and which are mutually accessible to one another, form a society of agents.

Agents which share the same communication conventions and which are mutually accessible to one another, form a society of agents. The complexity of the behavior of such a society depends in part upon the complexity of the conventions governing the semantic interpretation of the messages that agents can send one another.569To handle the combinatorial structures of interest to SFF, it will be sufficient toestablish that the SFF-family of digraphs can be correlated with societies of agents.First, every vertex is mapped onto to a pair (Λ, env), where Λ is an evaluatorand env is a mutable symbolic environment.

That is, vertices behave as iterative functions that are presented a time-sequence of inputs, and then in the context of the current state of env produce within a bounded time a time-sequence of outputs and possibly a revision of env.Second, every edge is associated with a transfer function which, given a time- sequence of values on the start vertex, yield a corresponding time-sequence of values on the stop vertex by applying the transfer function.Digraph automata are distributed networks of evaluators that communicate with each other by transfer functions. Among the expression that they may communicate are those that cause changes in env.It is possible to implement digraph automata as asynchronous networks but thisinteresting approach has been only partially implemented in [3]. There are also other execution styles.

There are also other execution styles. For example, since each node is connected to its neighbors in a precisely defined way, it is not unreasonable to adopt a computational approach in which the updates happen by a Jacobi sweep. That is, all nodes are evaluated at time t and all are determined values for time t+1, then the state is updated by swapping in the t+1 values. This style is used by cellular automata (see [12] for details).5. Concluding RemarksSuppose that a digraph automaton has been wired up and that each vertex hasbeen assigned a reasonable env. How would one use it to generate geometric realizations of 3-pseudomanifolds?First, if K is the 3-pseudomanifold being worked on, then there are huge numbersof realizations in Euclidean 3-space. But given the focus on a single component, this reduces to a specific set of requirements, specifications, and constraints which have been derived from the stratified product network that the component is being designed for.

It is a reasonable assumption that these requirements, specifications, and constraints can be coded in a symbolic environmentSecond, a boundary system (P,⊂ ) for K and the set of active patches in ∂K mustbe selected and symbol/value pairs for each of them must be added to the symbolic environment. The associated symbol will be the official way to reference a patch and the value will be the set of simplexes that are elements of the patch. Patches go proxy for geometric surfaces in 3-space that are homeomorphic to disks. These patches are the sources or sinks of edges which connect the part under study with other components of the product, thereby anchoring requirements, specifications, and constraints inherited from the product digraph, and thus marking the regions in which co-design will be required.Third, a reasonable sample of Morse slicings must be generated.

This will be a difficult task, but once a sample has been provided, it is exercise to code all of them in the symbolic environment.Fourth, produce a Reeb digraph for every Morse slicing in the sample. This canbe coded directly into our fictional digraph automaton, but with a surprising twist. Every570vertex and every edge of the Reeb digraph is assigned its own separate evaluator. Let’s call them v-evaluator and e-evaluators. The edges of the digraph automaton are connected in a manner prescribed by the Reeb digraph, that is, every edge in the Reeb digraph is mapped onto an edge path in the digraph automata which links up the source v- evaluator, the corresponding e-evaluator, and the sink v-evaluator. The local environment of each v-evaluator contains a symbol/value pair for the component of Morse slice paired with corresponding vertex in the Reeb digraph.

The local environment of each v-evaluator contains a symbol/value pair for the component of Morse slice paired with corresponding vertex in the Reeb digraph. The local environment of each e- evaluator has access to the local environment of its companions.To obtain geometric realizations for K at this point is a matter of outputtingpolygons from v-evaluators, blended cylindrical solids from e-evaluators, and surface geometries for the active patches – and of course to do all that while satisfying the product-based constraints!To make this agenda technically plausible will require a better understanding ofrole of algebraic topology in the optimal design of physical shapes and a working prototype of digraph automata.References[1]Adams, C.C. The Knot Book: An Elementary Introduction to the Mathematical Theory of Knots, W.H. Freeman, 2001. [2] Bott, R, “Lectures on Morse Theory, Old and New,” Bulletin (new series) AMS, vol 7 (1982); 331-358. [3] Boudreaux, J.C.

MBINED BUILD–TIME, ENERGY CONSUMPTION AND COST ESTIMATION FOR DIRECT METAL LASER SINTERINGM. Baumers*, C. Tuck*, R. Wildman*, I. Ashcroft*, E. Rosamond†, and R. Hague*Additive Manufacturing and 3D Printing Research Group, Faculty of Engineering, University of Nottingham, Nottingham, NG7 2RD, UK †Sustainability Research Group, Wolfson School of Mechanical and Manufacturing Engineering, Loughborough University, Loughborough, LE11 3TU, UK AbstractAs a single-step process, Additive Manufacturing (AM) affords full measurability with respect to process energy inputs and production cost. However, the parallel character of AM (allowing the contemporaneous production of multiple parts) poses a number of problems for the estimation of resource consumption. A novel combined estimator of build-time, energy consumption and production cost is presented for the EOSINT M270 Direct Metal Laser Sintering system.

A novel combined estimator of build-time, energy consumption and production cost is presented for the EOSINT M270 Direct Metal Laser Sintering system. It is demonstrated that the quantity and variety of parts demanded and the resulting ability to utilize the available machine capacity impact process efficiency, both in energy and in financial terms.IntroductionThe quantification of manufacturing cost informs the decision to adopt additive techniques in commercial applications. A number of cost estimators have been proposed for various additive technology variants (Alexander et al., 1998; Hopkinson and Dickens, 2003; Byun and Lee, 2006; Ruffo et al., 2006b; Wilson, 2006; Munguia, 2009).Moreover, a precise understanding of the emissions associated with manufacturing processes is essential regarding decision making towards sustainability.

The measurement of such emissions forms an important area of research in the field of industrial ecology, where a variety of methods are employed to analyze the interactions between human activity and the environment (Gößling-Reisemann, 2008). Particularly the quantification of carbon emissions, referred to as ‘carbon accounting’, requires a precise understanding of the energy flows associated with production processes (Vijayaraghavan and Dornfeld, 2010) and the emission characteristics of the local power grid (Jeswiet and Kara, 2008).No statement on the quality of a manufacturing process selection decision can be made without considering the manufacturing stage. For the private enterprise, the average manufacturing cost, normally measured per unit of output, is central.

For the private enterprise, the average manufacturing cost, normally measured per unit of output, is central. Considering the wider social implications of energy consumption, manufacturing process energy consumption is also becoming increasingly important (Taylor, 2008).The supply chains found in modern manufacturing are often complex and long (Foran et al., 2005). Industrial trends towards globalization, concentration on core activities, shorter product lifecycles and the increasing focus on customer needs will add to supply chain complexity (Blecker and Kersten, 2006). The resulting opacity poses a significant barrier to the measurement and minimization of resource consumption in the manufacturing stage of the product life cycle. The single-step nature of additive processes affords full measurability with respect to process energy inputs and production costs.

The single-step nature of additive processes affords full measurability with respect to process energy inputs and production costs. This is especially interesting as additive processes are able to efficiently create complex product geometries (Hague et al., 2004; Rosen, 2007).932The Additive Manufacturing (AM) process allows the contemporaneous production of multiple, potentially unrelated, components. The technology has therefore been described as a ‘parallel’ manufacturing technology (Ruffo et al., 2006b). However, this aspect creates two problems for the modelling and optimization of additive production:1.To arrive at some summary measure of cost or energy consumption per part, it is necessary to attribute the total cost and energy consumption incurred during each build.

In the single product case this can be done by dividing the total cost and energy consumption of the build by the number of parts contained.2.It has been shown that the degree of capacity utilization affects the resulting metrics of process efficiency (Ruffo et al., 2006b; Ruffo and Hague, 2007; Baumers et al., 2011). Therefore, to be able to claim that a build is performed at minimum cost and energy consumption, the available machine capacity should be fully utilized. In this research, the issue is approached by implementing a packing algorithm that fills the available capacity with parts in an optimized configuration.A major additive technology variant used to manufacture metal components is Direct Metal Laser Sintering (DMLS), belonging to the category of powder bed fusion processes (ASTM, 2012). This paper analyzes the EOSINT M270 platform (EOS GmbH, 2010), belonging to the class of DMLS technology.

This paper analyzes the EOSINT M270 platform (EOS GmbH, 2010), belonging to the class of DMLS technology. DMLS operates as follows: a three-dimensional representation of the product geometry is digitally cut into discrete slices. These slices are then transmitted to the DMLS machine, which recombines them in a layer-by-layer sequence. To this end, the EOSINT M270 selectively scans the surface of a metal powder bed with a 200 W fiber laser, effectively creating a thin, planar slice of solid part geometry. Once the sintering of the layer is complete, a fresh 0.02 mm increment of metal powder (in this case stainless steel) is deposited and the sintering of the next layer commences. This cycle is repeated until the build is complete.The proposed model of DMLS build-time, energy consumption and cost suggests that average production cost and energy consumption should not be viewed as dependent on production quantity.

This notion has been expressed implicitly for the other additive technology variants (Hopkinson and Dickens, 2003). Rather than being determined by production quantity, the current article suggests that the technology user’s ability to fill the available build space is the prime determinant of efficient technology operation. This conforms to the observation that manufacturers will ideally include as many parts as possible in individual builds (Ruffo and Hague, 2007).MethodologyBuild time estimation forms the basis for several AM production cost models in the literature (Alexander et al., 1998; Byun and Lee, 2006; Campbell et al., 2008). Ruffo et al. (2006b) propose an AM costing model viewing the total cost of a build, CBuild, as the sum of all direct raw material costs and the indirect costs of operating the machine. The activity based costing technique (Atrill and McLaney, 1999) is a suitable costing method for capital- heavy production processes such as AM.

The activity based costing technique (Atrill and McLaney, 1999) is a suitable costing method for capital- heavy production processes such as AM. CBuild can thus be modeled as:̇(1)933The direct costs are obtained by multiplying the mass of deposited material (mMaterial) by the cost of the raw material (CMaterial) per kilogram (£/kg, where a $/£ exchange rate of 1.56 is used). The indirect costs are calculated by multiplying the total build time, TBuild, by an indirect cost rate, ĊIndirect (£/s). For an estimate of cost per part, CBuild is divided by the number of parts contained in the build.Therefore, this method is only suitable for builds containing multiple instances of the same part. As a premise, this is quite alien to the idea of AM being used to flexibly build different parts in parallel (Ruffo and Hague, 2007).

As a premise, this is quite alien to the idea of AM being used to flexibly build different parts in parallel (Ruffo and Hague, 2007). By dividing the cost estimate of the build (CBuild) by the number of parts contained in the build, this type of model is also used to describe a relationship between production quantity and average part cost. Where production quantities are not sufficient to fill the available build space, the model assumes that the available capacity remains empty.However, it has been observed in practice that additive technology users fill the available build space with as many parts as possible (Ruffo and Hague, 2007). Should the demand for parts not be sufficient to fill the workspace, the technology user has the option to sell excess machine capacity to external bidders.

Should the demand for parts not be sufficient to fill the workspace, the technology user has the option to sell excess machine capacity to external bidders. If the problem of excess capacity persists it would further be possible to reduce manufacturing cost by switching to an additive platform with a smaller capacity.For models of manufacturing cost (especially when taking the form of cost functions) it is normally considered necessary that only efficient technology usage is taken into account (Else and Curwen, 1990). Research has shown that some additive systems operate efficiently only where the available capacity is fully utilized (Ruffo et al., 2006b; Ruffo and Hague, 2007; Baumers et al., 2011). However, to be able to claim that the additive process is used at full capacity, it is central that the build is configured with the maximum number of parts in the build.

However, to be able to claim that the additive process is used at full capacity, it is central that the build is configured with the maximum number of parts in the build. To approach the problem of filling build volumes, workspace packing algorithms have been developed (Wodziak et al., 1994; Nyaluke et al., 1996; Ikonen et al., 1997). For an efficient packing outcome, Hur et al. (2001) suggest using voxel approximations of part geometry, allowing nested configurations of parts.To solve the issue of efficient technology usage in energy consumption and cost estimation, this research bases the estimator on an automated build volume packing technique. This functionality is designed to arrange parts in the build volume such that the resulting build configuration is optimized.

This functionality is designed to arrange parts in the build volume such that the resulting build configuration is optimized. To allow a relatively simple implementation, a number of decisions were made in the design of the algorithm:To achieve a dense packing result, the algorithm is based upon rough voxel representations of the analyzed parts (with a resolution of 5 mm). This effectively discretizes the problem of placing irregular and continuous geometries. Figure 1 illustrates the conversion of continuous part geometry into a voxel approximation. To avoid problems of anisotropic material properties occurring in the DMLS process, part rotation is constrained to the vertical axis. Rotation is also limited to discrete 90° steps.The parts are inserted into the build volume in a fixed sequence which is predetermined. Further, the algorithm is able to insert a variety of parts.

Rotation is also limited to discrete 90° steps.The parts are inserted into the build volume in a fixed sequence which is predetermined. Further, the algorithm is able to insert a variety of parts. It ensures that at least one instance of each assessed part type is included in the build volume.934As an evaluation criterion, the algorithm uses a ‘barycentric’ packing heuristic. This means that the algorithm computes the combined center of mass of all parts contained in the build volume. When a further part is inserted, it is moved such that its center of mass is as close as possible to the existing combined center of mass. This produces a dense configuration of parts and also allows for nesting of components.In the DMLS process, all parts must be connected to a removable build plate, forming the build volume floor. This is done to prevent part deformation due the heat introduced during the process and to allow heat dissipation into the machine frame.

This is done to prevent part deformation due the heat introduced during the process and to allow heat dissipation into the machine frame. Thus, the algorithm only considers arrangements in which all parts are placed on the substrate, effectively limiting part movement to the X/Y plane.Figure 1: Conversion of a test part into a voxel approximationAfter implementing the required build volume packing functionality, the next step in the construction of the combined estimator is to decide on a basket of test parts. The composition of this basket is chosen to be representative of the products commercially manufactured using DMLS and to reflect variation in product size, geometry, and application. The used basket of parts is shown in Figure 2.Figure 2: A basket of representative parts935In an attempt to emulate a realistic application of DMLS, the workspace is populated with parts drawn from this basket for a full build experiment.

Two further build experiments were performed to validate the model, each holding a single part from the basket. Build progress and process energy consumption were monitored throughout the build experiments with a Yokogawa CW240 digital power meter configured to a time resolution of 1s. Thermal management considerations did not flow into the build configurations.To generate the build time and energy consumption data required for the combined estimator, a further build experiment was performed and monitored with a higher time resolution (100 ms). In this experiment, a single specifically designed test part was built directly on the build platform. As shown in Figure 3, this test part features a layered design to give an insight into the EOSINT M270’s material deposition process.Figure 3: Design of the power monitoring test partNext to actual process energy consumption, additional energy is consumed for the separation of the parts from the removable build plate.

This is done in an ancillary wire erosion process using an Agie Charmilles CUT20 wire eroder. This setup was also power monitored.Once the build configuration is determined through an execution of the build volume packing algorithm and the data on build speed, energy consumption and production cost have been collected, the combined model can be constructed. In order to do this, the first step is to estimate build time, TBuild, which is obtained by combining data from a hierarchy of elements of time consumption:fixed time consumption per build operation, TJob, including, for example, machine atmosphere generation and machine warm-up,total layer dependent time consumption, obtained by multiplying the fixed time consumption per layer, TLayer, by the total number of build layers l,the total build time needed for the deposition of part geometry approximated by the voxels.

The triple Σ operator in equation (2) expresses the summation of the time needed to process each voxel, TVoxel xyz, in a three-dimensional array representing the discretized build configuration:936( ) ∑ ∑ ∑(2)No allowance is made for build preparation and machine cleaning. It is felt that the time spent on these activities is difficult to measure and very much at the discretion of the machine operator. It could be argued that these activities take place during the non- operational hours.Total energy investment, EBuild, can be modeled similarly to equation (2). However, a purely time-dependent element of power consumption must be expected in the continuous operation of the AM machine. This is denoted by the energy consumption rate ĖTime (measured in MJ/s), which is multiplied by TBuild to estimate total time-dependent energy consumption.

This is denoted by the energy consumption rate ĖTime (measured in MJ/s), which is multiplied by TBuild to estimate total time-dependent energy consumption. Modelling ĖTime as a constant reflects the mean baseline level of energy consumption throughout the build, originating from continuously operating machine components such as cooling fans, pumps, and the control system.EJob contains all energy consumption attributable to the build job, including energy consumed by the wire erosion process to harvest the parts from the build plate. Analogous to build time estimation, ELayer denotes fixed elements of energy consumption per build and layer, for a total number of layers, l. Further, the geometry-dependent energy consumption is obtained by adding all energy consumption associated with actual material deposition, EVoxel xyz , throughout the discretized workspace. Please note that EVoxel xyz does not contain time-dependent power consumption.

Please note that EVoxel xyz does not contain time-dependent power consumption. The empirical data on EVoxel xyz were obtained by monitoring machine energy consumption during the scanning process and then subtracting the energy associated with the energy consumption rate ĖTime, ensuring that this element of energy consumption is not counted twice. Thus, EBuild can be modeled as follows:( ̇ ) ( ) ∑ ∑ ∑(3)This energy consumption model should not be interpreted as showing how total AM energy consumption can be attributed to individual subunits of the EOSINT M270 platform. The specification was chosen with the goal of implementing a voxel-based energy consumption estimator. Moreover, both the time and energy estimators possess additional information on the real Z-height of the parts contained in the build.

Moreover, both the time and energy estimators possess additional information on the real Z-height of the parts contained in the build. This is done to avoid large estimation errors arising from the inclusion of empty layers.After developing the build time and energy consumption techniques, the next step towards the combined estimator is the construction of an activity-based cost (ABC) estimator of the type devised by Ruffo et al. (2006b). The cost estimate for the build, CBuild, is computed from data on the total indirect costs and direct costs incurred. All data used in the costing model are summarized in Table 1. The current research estimates the total indirect cost rate of operating the EOSINT M270 at £26.64 per hour. It is noteworthy that the system incorporates an N2 generator, hence no protective gas from external sources is needed.937Table 1: Cost model elements (adapted from Ruffo et al.

Total raw material costs are calculated by multiplying the total mass w of all parts included in the build (including support structures) with the price per kilogram of the stainless steel 17-4 PH powder, PriceRaw material (78.81 £/kg). Thus, any raw material losses are ignored. The expenditure for energy enters the model by multiplication of the energy consumption estimate, EBuild, with the mean price of electricity for the manufacturing sector in the UK, PriceEnergy, currently around 0.018 £/MJ (according to DECC, 2010). The total cost estimate for the build, CBuild, can be expressed as:( ̇ ) ( ) ( )(4)Results and DiscussionA full-capacity build experiment is configured by executing the build volume packing algorithm. The resulting full build configuration is shown in Figure 4. Of the available 2,025 build volume floor voxels, 92.6% were occupied. A total of 85 parts were inserted, utilizing 19.78% of the used build volume cuboid (225 mm x 225 mm x 52 mm).

Of the available 2,025 build volume floor voxels, 92.6% were occupied. A total of 85 parts were inserted, utilizing 19.78% of the used build volume cuboid (225 mm x 225 mm x 52 mm). This value includes the auxiliary structures needed to anchor the overhanging part geometry on the substrate.938Figure 4: Full build configuration of basket partsThe full build experiment (including the wire erosion process) consumed a total of 1059.56 MJ of energy. Using the cost model specified in equation (4), CBuild is estimated at £3,218.87. Individual part cost and energy usage are identified through their share of total product mass (4.167 kg).

Using the cost model specified in equation (4), CBuild is estimated at £3,218.87. Individual part cost and energy usage are identified through their share of total product mass (4.167 kg). A summary of the parts produced in the full build and estimates of energy usage and production cost are presented in Figure 5:Figure 5: Estimates of energy consumption and production cost per partThe final specifications of the time and energy estimators, equations (2) and (3), are obtained from a least squares regression of the time and energy consumption data recorded during the experiment containing the layered power monitoring test part, as shown in Figure 3.The obtained parameters αTime (10.82 s) and αEnergy (0.008 MJ) are multiplied by the number of layers in the build l in order to obtain layer dependent time and energy consumption.

The parameters expressing the time and energy attributable to the scanning of 1939mm² during the build, βTime (0.0125 s) and βEnergy (0.000013 MJ), are then used in conjunction with the layer thickness lt (0.02 mm) and a measure of occupancy of each voxel to calculate total time and energy consumption per voxel, TVoxel xyz and EVoxel xyz. The rate of occupancy ROi in each voxel is modeled as the ratio of the volume of part i occupying this voxel (VPi) and the volume of the voxel approximation for part i (VAi):(5)Thus, for each (5 mm)³ voxel in the position xyz holding 250 (= 5 mm / lt) layers andcontaining part i, the build time and energy consumption can be approximated:(6)(7)This is combined with an estimated fixed time and energy consumption for machine start-up TJob (63 s) and EJob (142.58 MJ, including wire erosion).

The start up process is very rapid on this system as no warm up is required and the build chamber is continuously flooded with N2 during build activity. Time dependent power consumption is obtained by multiplying the base line energy consumption rate ĖTime (0.0015 MJ per s) with TBuild. Thus, the estimates of TBuild and EBuild are obtained as follows:( ) ∑ ∑ ∑(8)( ) ( ) ∑ ∑ ∑(9)The time and energy consumption model specified in equations (8) and (9) is experimentally validated. This is done by comparing the calculated estimates to the measured time and energy consumption during the three build experiments containing parts from the representative basket (Figure 2). Validation is performed for the full build at maximum machine capacity (shown in Figure 4) and two single part builds, the bearing block and the turbine wheel.The results of the validation experiments and the corresponding estimates of TBuild and EBuild are presented in Table 2.

Note that the validation does not include the energy consumed by the ancillary wire erosion process. It should also be mentioned that some of the venturi parts had an incorrect orientation during the build, which led to build failure for the affected parts in the final stages of the full build. However, this was deemed to have had a negligible effect on the presented results.940Table 2: Confronting the estimates with experimental resultsExperimentTime consumedModel estimate TBuildErrorEnergy usageModel estimate EBuildErrorFull Build experiment388031 s354806 s8.56 %917.10 MJ879.93 MJ4.05 %Single Bearing block93302 s92338 s1.03 %215.48 MJ223.13 MJ3.55 %Single Turbine wheel31224 s28504 s8.71 %72.73 MJ66.80 MJ8.15 %The observed errors are likely to originate from the use of an idealized test part (Figure 3) in the experiment that provided the data. Compared to other build time estimators (Campbell et al.

Compared to other build time estimators (Campbell et al. 2008; Munguia 2009; Ruffo et al. 2006a; Wilson 2006) the errors reported in Table 2 indicate that the developed time estimation functionality performs robustly.This research thus demonstrates that the impacts of the fixed process elements in the DMLS process, which may be job dependent (such as machine start-up) or layer dependent (such as powder re-coating), are amortized over the number of parts contained in the build. This has been previously suggested by Baumers et al (2011) for energy consumption.The presented technique appears appropriate for additive processes as it can be used to estimate specific energy consumption and production cost for various build configurations with multiple types of parts. To test the effect of various build compositions, eight different configurations with varying degrees of capacity utilization were estimated. The results for process energy consumption range from 1.96 MJ/cm³ to 3.61 MJ/cm³.

The results for process energy consumption range from 1.96 MJ/cm³ to 3.61 MJ/cm³. Assuming a material density of 7.80 g/cm³, this corresponds to specific energy consumption ranging from 251.28 MJ/kg to 462.82 MJ/kg. This is slightly higher than reported in previous research, ranging from 241 MJ/kg to 339 MJ/kg (Baumers et al., 2011). This difference is likely to originate from the inclusion of the energy consumed by the wire erosion process. In terms of production cost, the eight build configurations led to results ranging from 5.71 £/cm³ to 7.44 £/cm³.The full build configuration shown in Figure 4 resulted in the lowest estimated energy consumption and production cost on the EOSINT M270 (1.96 MJ/cm³ and 5.71 £/cm³). This underlines the statement made in the introduction that for cost and energy consumption metrics to reflect efficient machine operation it is important to consider full capacity utilization.

This underlines the statement made in the introduction that for cost and energy consumption metrics to reflect efficient machine operation it is important to consider full capacity utilization. The results thus show that the configuration with the highest packing density is likely to lead to the most efficient build. This points to the conclusion that the user’s ability to fully utilize the available build space is an important determinant of DMLS cost and energy consumption.941ConclusionsThis research demonstrates the construction of a combined estimator of build-time, energy consumption and cost for parallel additive techniques such as DMLS which reflects technically efficient machine operation. The application of this methodology shows that the cost and energy consumption of the DMLS process are determined by the user’s ability to fill the available build space.

The application of this methodology shows that the cost and energy consumption of the DMLS process are determined by the user’s ability to fill the available build space. Further, the proposed method can be used to estimate the production of own designs in build volumes that are populated (where necessary) with parts drawn from a representative basket.The developed methodology has been applied to DMLS, which is a laser-based additive platform employing a powder bed. While the results are likely to be extensible to later generations of DMLS systems (such as the EOSINT M280), it is unclear whether they are applicable to other additive processes. These could be platforms operating with a powder bed (for example, Electron Beam Melting) or those with an entirely different operating principle (for example, Fused Deposition Modelling). Further research is needed in this area.

Further research is needed in this area. Moreover, the model described in this paper is limited to so called ‘well structured’ costs of manufacturing (Son, 1991). Ill structured costs arising from factors such as build failure, machine idleness and inventory expenses are ignored.References1. Alexander, P., Allen, S., and Dutta, D., 1998. Part orientation and build cost determination in layered manufacturing. Computer-Aided Design, 30 (5), pp.343-356.2. ASTM, 2012. Standard F2792 - 12a, 2012, Standard Terminology for Additive Manufacturing Technologies. ASTM International, West Conshohocken, PA, 2003, DOI: 10.1520/C0033-03, www.astm.org.3. Atrill, P., and McLaney, E., 1999. Management Accounting for Decision Makers. 2nd ed. London: Prentice Hall Europe.4. Baumers, M., Tuck, C., Wildman, R., Ashcroft, I., and Hague, R., 2011c. Energy inputs to additive manufacturing: does capacity utilization matter? Proceedings of the Solid Freeform Fabrication (SFF) Symposium.

MBINED EFFECTS OF MONTMORILLONITE CLAY, CARBON NANOFIBER, AND FIRE RETARDANT ON MECHANICAL AND FLAMMABILITY PROPERTIES OF POLYAMIDE 11 NANOCOMPOSITESBlake Johnson, Eric Allcorn, Min G. Baek, and Joseph H. Koo* The University of Texas at Austin, Dept. of Mechanical Engineering, 1 University Station, C2200, Austin, TX 78712-0292, USAandCarla Leer Lake and Patrick Lake Applied Sciences, Inc., 141 W. Xenia Ave, Cedarville, OH 45314, USACorresponding author: jkoo@mail.utexas.eduABSTRACTThis paper is focused on the development of polyamide 11 (PA11) nanocomposites with enhanced fire retardant (FR) properties for application in selective laser sintering (SLS). Test specimens of PA11 containing various percentages of intumescent FR additive, montmorillonite (MMT) clay, and carbon nanofiber (CNF) were prepared via the twin screw extrusion technique.

Test specimens of PA11 containing various percentages of intumescent FR additive, montmorillonite (MMT) clay, and carbon nanofiber (CNF) were prepared via the twin screw extrusion technique. The combined effects of MMT clay, CNF, FR additives on the mechanical and flammability properties of these PA11 nanocomposites are studied. Izod impact testing, tensile testing, and SEM analysis of are used to characterize mechanical properties. UL-94 and SEM analysis of char surfaces are used to characterize the flammability properties of these materials. Results are analyzed to determine any synergistic effects among the additives to the material properties of PA11.1. INTRODUCTIONSelective Laser Sintering (SLS) is a method of rapid manufacturing (RM) that involves the fabrication of parts through the building up of layers of a polymer powder.

INTRODUCTIONSelective Laser Sintering (SLS) is a method of rapid manufacturing (RM) that involves the fabrication of parts through the building up of layers of a polymer powder. Each layer is very thin (0.076-0.152mm) and is added to the part when a computer-controlled laser heats and fuses the powder particles into a solid material, which is also fused to the solid layer that was previously fused below it (schematic is shown in Figure 1). This process allows for a part to be manufactured directly from a 3D CAD file. The 3D design is sliced into layers that are then fused together using an SLS machine. The SLS technique has developed significantly over the past several years and is used in the manufacturing of a modest variety of commercial and industrial parts and prototypes [1].Figure 1: Schematic of SLS process.Current polymers that are widely used in SLS are polyamide (PA) 11, PA12, and glass-filled PA12.

Commercial manufacturing endeavors have been undertaken that make use of PA11 and PA12, such as Boeing’s use of PA11 SLS manufacturing to produce air ducting systems for F18 Fighter jets [2] and Siemen’s use of PA12 SLS manufacturing to produce hearing aid shells. Despite the successes in the application of these SLS manufactured PA11 and PA12 parts, these materials have number of poor characteristics that limit their applications. Both PA11 and PA12 have relatively poor thermal, flammability, and mechanical properties when compared to other commercially available thermoplastics. They are limited to a use temperature of roughly 71°C, with brief exposure to temperatures up to 121°C. In addition, both of these materials are quite flammable.

They are limited to a use temperature of roughly 71°C, with brief exposure to temperatures up to 121°C. In addition, both of these materials are quite flammable. The tensile strength and moduli of these PA polymers are also mediocre, especially when compared to traditionally manufactured polymers such as PEEK [3].It is possible through the use of additives to modify and improve select properties of a polymer to better perform its designed function. Such additives, both micron-sized and nano-sized, have been successfully implemented to improve the flammability and select mechanical properties of PA11 [3-6]. In conventional methods, FR additives like inorganic metal-oxides or halogens are used to improve the flammability characteristics of polymers. In order achieve desirable FR properties the weight percentage of FR additives in these polymer mixtures is typically greater than 30%.

In order achieve desirable FR properties the weight percentage of FR additives in these polymer mixtures is typically greater than 30%. However, the inclusion of FR additives also results in the degradation of the polymer’s mechanical properties [7].Nanomaterials have also been successfully used as additives to improve the properties of polymers. Nanomaterials are defined as materials with structured components that have at least one dimension in the nanometer-scale (10-9 meter). There are numerous varieties of nanomaterials being implemented into different polymers to improve the flammability of polymer materials: nanoclay, carbon nanofibers (CNFs), or multi-walled carbon nanotubes (MWNTs). These different materials have differing degrees of effectiveness, however they all offer material improvements at much lower weight percentages (<7wt%).

These different materials have differing degrees of effectiveness, however they all offer material improvements at much lower weight percentages (<7wt%). In addition, the inclusion of nanomaterials typically results in improved mechanical properties as well [3,4].While these additives are helpful when used separately, they are often most beneficial when used in combination with each other if they are compatible. It has been shown that combining conventional FR additives with nanomaterial additives can greatly increase the effectiveness of both, as well as reduce or reverse the negative effect of conventional FR on the mechanical properties of polymers [1].

In addition, two nanomaterial additives used at once have the potential to improve various polymer properties greater than the same amount of either nanomaterial would individually [5,6].The purpose of this paper is to study the combined effects of a conventional FR, montmorillonite (MMT) nanoclay, and CNF into a PA11 polymer base. Focus will be placed on the mechanical and flammability properties of the resulting composites in comparison to neat PA11 and PA11 composites lacking the synergistic effects of multiple additives.2. EXPERIMENTATION2.1 Materials2.1.1 Polymer The base polymer used for this study is Rilsan® PCG LV polyamide 11 (PA11) manufactured by Arkema. PA11 is a versatile thermoplastic polymer noted for its positive properties of abrasion, chemical, and heat resistance, good resistance to creep, and a low coefficient of friction.

PA11 is a versatile thermoplastic polymer noted for its positive properties of abrasion, chemical, and heat resistance, good resistance to creep, and a low coefficient of friction. PA11 is an intermediate melting material with a melting temperature of less than 200°C.2.1.2 Fire Retardant Additive Clariant manufactures different FR additives known as Exolit® OP grades for various types of plastics. The specific FR used in this study is OP1230. It is an intumescent FR additive based on metal phosphinate. It is a white powder containing phosphorus and other propriety FR. It is non-halogenated, hydrophobic, and exhibits good thermal stability. It was developed as a general purpose FR additive for plastics.2.1.3 Nanoclay The nanoclay used in this study is Cloisite® 30B manufactured by Southern Clay Products.

It was developed as a general purpose FR additive for plastics.2.1.3 Nanoclay The nanoclay used in this study is Cloisite® 30B manufactured by Southern Clay Products. It is a natural montmorillonite clay modified with an organic modifier.2.1.4 Carbon Nanofibers Carbon nanofibers are discontinuous graphite filaments with diameters less than 200nm and varying lengths manufactured by Applied Sciences, Inc. The specific grade used in this study is PR24XT-PS CNF.2.1.5 Processing The materials used in this experiment were processed via co-rotating twin- screw extrusion, a melt blending process. Masterbatches of PA11/10% MMT and PA11/10% CNF were produced prior to the creation of the final PA11/MMT/CNF/FR blends. The specific operating conditions for the melt-blending process are shown in Tables 1-3.

It measures the energy required to fracture a notched specimen at relatively high rate bending conditions. A pendulum with adjustable weight is released from a known height; a rounded point on the tip of the pendulum makes contact with a notched specimen [8]. The specific device used in our testing is the Shanta Engineering Izod/Charpy Impact Tester. The specimen layout and Izod tester is shown in Figure 2.Figure 2: Shanta Engineering Izod/Charpy Impact Tester (left) and layout of Izod test specimen according to ASTM D256 [9].2.2.2 Tensile Tests Tensile tests are performed for several reasons and the results of tensile tests are frequently included in material specifications to ensure quality. Tensile testing involves locking a specimen into a testing setup and applying a tensile load to specimen until it breaks. The data recorded through this testing can be used to determine numerous material properties [10].

The data recorded through this testing can be used to determine numerous material properties [10]. In our application we are interested in the ultimate tensile strength, elongation at break, and tensile modulus. The specific device used in our testing is an Instron Tensile Tester.2.2.3 Microstructural Analysis In order to better understand the microstructural mechanisms behind our mechanical test results, we will analyze the fracture surface of our post-test Izod and tensile specimens. This analysis will give us a greater qualitative understanding of how our various additives affect the structure and properties of our composite materials. Similarly, microstructural analysis will be conducted on burnt UL-94 specimens to understand the protective mechanisms.

Similarly, microstructural analysis will be conducted on burnt UL-94 specimens to understand the protective mechanisms. We used the Quanta 650 SEM/ESEM to perform fractural and charred surfaces analysis of our materials.2.2.4 UL-94 Vertical Burn Test The test for flammability of plastic material used in parts for electronic devices and appliances is commonly known as UL-94. It is a flammability specification issued by the Underwriters Laboratories Inc [11].The flammability specification UL-94 contains several small-scale flame tests procedures that define parameters for flame testing of plastics. In this specification, the behavior of plastic when exposed to fire is expressed quantitatively. The amount of material burned refers to the length of sample that burns in a specified period of time.

In this specification, the behavior of plastic when exposed to fire is expressed quantitatively. The amount of material burned refers to the length of sample that burns in a specified period of time. We performed the vertical burn test, in which a flame is applied to the base of the specimen held in the vertical position and the extinguishing times are determined upon removal of the test flame, as well as whether the specimen produces flaming drips (see Figure 3 for test layout). Ratings for UL-94 range from V-2, which is the worst rating for a material that still passes, to V-0, which is the best rating [11].Figure 3: Testing layout for the UL-94 vertical burn test.3. RESULTS3.1 Mechanical Properties3.1.1 Izod Impact Strength The results from Izod impact testing of our composites are shown in Table 4. Formulation 2, which includes 3.5% MMT, 3.5% CNF, and 0% FR, has the highest Izod impact strength.

Formulation 2, which includes 3.5% MMT, 3.5% CNF, and 0% FR, has the highest Izod impact strength. Formulation 6, which includes 3.5% MMT, 3.5% CNF, and 20% FR, has the lowest Izod impact strength. There are no significant differences between 3.5% MMT, 3.5% CNF and 2.5% MMT, 2.5% CNF. However, greater FR percentages tend to lead to lower Izod impact strength, as shown in Figures 4 and 5.

There are no significant differences between 3.5% MMT, 3.5% CNF and 2.5% MMT, 2.5% CNF. However, greater FR percentages tend to lead to lower Izod impact strength, as shown in Figures 4 and 5. The Izod impact strength of neat PA11 is remarkably high compared to the polymer blends, and for most formulations the decrease from neat PA11 is greater than 90%.Figure 4: Izod impact strength for PA11/2.5% MMT/2.5% CNF composite as a function of weight percent FR OP1230.Figure 5: Izod impact strength for PA11/3.5% MMT/3.5% CNF composite as a function of weight percent FR OP1230.Table 4: Izod Impact Strength of PA11 CompositesFormulation MMT CNF FRIzod Impact Strength (J/m)Neat PA11 #1 #3 #4 #2 #5 #60% 0% 2.5% 2.5% 0% 2.5% 2.5% 15% 2.5% 2.5% 20% 3.5% 3.5% 0% 3.5% 3.5% 15% 3.5% 3.5% 20%0%113.7 12.7 6.6 11.3 18.4 6.7 5.73.1.2 Tensile Testing The results for tensile testing of our composites are shown in Table 5.

Figure 6 shows tensile strength of PA11 composites as a function of weight percent FR OP1230. Polymer blends which include 2.5% MMT and 2.5% CNF have higher tensile strength but the differences between them are not considerable. The weight percentage of FR also seems to have minimal effect on the tensile strength of the materials. However, all materials have a lower tensile strength than neat PA11.Figure 6: Tensile strength for PA11 composites as a function of weight percent FR OP1230.Figure 7 shows that polymer blends with 3.5% MMT and 3.5% CNF generally have higher tensile modulus than those with only 2.5% MMT and 2.5% CNF. Formulation 5, which includes 3.5% MMT, 3.5% CNF, and 15% FR, has the highest tensile modulus. Formulation 3 which includes 2.5% MMT, 2.5% CNF, and 15% FR has the lowest tensile modulus. The weight percentage of FR OP1230 also seems to increase the tensile modulus, though the effect is not very significant.

The weight percentage of FR OP1230 also seems to increase the tensile modulus, though the effect is not very significant. Compared to neat PA11, the polymer blends are more brittle but the tensile modulus is noticeably increased.Figure 7: Tensile modulus for PA11 composites as a function of weight percent FR OP1230.Figure 8 shows that formulation 3, which includes 2.5% MMT, 2.5% CNF, and 15% FR, has the highest elongation at break. Formulation 2, which includes 3.5% MMT 3.5% CNF and 0% FR,has the lowest elongation at break. The materials with higher concentrations of MMT and CNF tend to be more brittle and have lower elongation at break. There are mixed results on the effect of FR weight percent. For the 2.5% MMT 2.5% CNF compounds, FR increases elongation slightly, but for 3.5% MMT 3.5% CNF there is no significant change as a result of FR inclusion.

For the 2.5% MMT 2.5% CNF compounds, FR increases elongation slightly, but for 3.5% MMT 3.5% CNF there is no significant change as a result of FR inclusion. All of the composites, however, show a significant decrease in elongation at break compared to neat PA11, which has more than 200 times higher elongation at break. Such a dramatic decrease in elongation is typical of nanocomposite materials.Figure 8: Elongation at break for PA11 composites as a function of weight percent FR OP1230.Table 5: Tensile Testing Results for PA11 CompositesFormulation MMT CNF FR 0% 0% 0% Neat PA11 2.5% 2.5% 0% #1 2.5% 2.5% 15% #3 2.5% 2.5% 20% #4 3.5% 3.5% 0% #2 3.5% 3.5% 15% #5 3.5% 3.5% 20% #6Tensile Strength (MPa) 41.8 33.2 32.1 32.3 30.1 30.7 32.0Elongation at Break 234% 1.74% 3.64% 2.91% 0.7% 0.91% 0.78%Tensile Modulus (MPa) 901 1931 1917 2209 2323 2453 24493.1.3 Fractural Surface SEM After completion of Izod impact testing, SEM images were taken of the fractural surface of the samples.

The polymer material is insulating so the samples had to be coated prior to SEM analysis. Representative images from each formulation are shown in Figures 8-10 below. In Figure 8, the two materials are shown to have a relatively smooth fracturesurface with strands of CNF protruding out from the surface. The second formulation has a higher concentration of CNF than the first (3.5% vs. 2.5%), which can be seen by the greater number of visible CNFs in the SEM image. The MMT clay is too small to be observed by the SEM used in this experiment. The protrusion of the CNFs from the fracture surface would seem to suggest that CNF would reinforce the material structure against fracture, but from the Izod results we know that these composites have much lower Izod resistance than the neat PA11. We have no definitive reason for this result based on the SEM, but it could be due to an effect from the MMT clay, which we are unable to observe.

We have no definitive reason for this result based on the SEM, but it could be due to an effect from the MMT clay, which we are unable to observe. This could be due to the weak interfacial bonding between the CNF and PA11 polymer.In Figure 9, the fracture surface of formulations 3 and 4 are shown to be quite rough when compared to those of the first two formulations. This could be due to the inclusion of micro-scale FR particles, which act as defects or weak points in the material structure that help to propagate the fracture. The FR particles themselves do not fracture, though, so they produce rough irregularities on the fracture surface. The Izod results agree with this analysis as the Izod resistance tends to decrease with the inclusion of FR. The protrusion of CNF can still be seen on the surface as well.In Figure 10, the fracture surface of formulations 5 and 6 are shown to be similar to those shown in Figure 9, but with much greater CNF protrusion.

This is to be expected as the weight percentage of CNF is higher in these formulations. The rest of the analysis is similar to that of Figure 9, where the FR particles create a rough surface and lower Izod resistance.Figure 9: SEM images of the Izod fracture surface for formulations 1 (left) and 2 (right).Figure 10: SEM images of the Izod fracture surface for formulations 3 (left) and 4 (right).Figure 11: SEM images of the Izod fracture surface for formulations 5 (left) and 6 (right).Flammability Properties3.2.1 UL-94 Results The results for the UL-94 vertical burn test of our materials are shown in Table 6. The baseline PA11 material received a V-2 rating because it produce burning drips, but was still self-extinguishing in less than 30 seconds. The two formulations with MMT clay and CNF, but no FR additive failed the test. In the PA11/2.5% MMT/2.5% CNF formulation, the material burned quickly once ignited and produced numerous viscous drips.

In the PA11/2.5% MMT/2.5% CNF formulation, the material burned quickly once ignited and produced numerous viscous drips. The PA11/3.5% MMT/3.5% CNF also burned readily once it was ignited, but it only produced drips on 6 out of 10 tests and the drips were very viscous. All of the formulations with FR additive passed the test with V-0 ratings. They all immediately self-extinguished and produced no drips during testing. Photos of the post-test specimens are show in Figure 7.Table 6: UL-94 Vertical Burn Ratings for PA11 CompositesFormulationMMT CNFFRUL-94 RatingNeat PA110%0%0%V-2#1 #3 #42.5% 2.5% 0% 2.5% 2.5% 15% 2.5% 2.5% 20%Fail V-0 V-0#2 #5 #63.5% 3.5% 0% 3.5% 3.5% 15% 3.5% 3.5% 20%Fail V-0 V-0The results for the formulations without FR additive are as expected.

It is not unusual for nanomaterial additives alone to fail to improve a material’s flammability rating and there are typically most beneficial when used to improve the effectiveness of conventional FR additives [3]. The results for the remaining formulations are quite promising. In previous tests, where MMT clay and CNF were combined with intumescent FR additives separately, such significant improvements in flammability properties were not observed with such low percentages of FR. In those experiments there were no formulations that were able to reach V-0 rating with only 15% FR, and there were several formulations that failed to reach it even at 20% FR [1]. These results suggest that there is a synergistic effect between MMT clay and CNF in improving the flammability of PA11 composites when combined with conventional FR additives.Figure 12: Post-test photos of UL-94 test specimens.

Formulations 1 and 2 (left-most two photos) can be seen to have burned significantly while formulations 3-6 (in ascending order from left to right) show no significant burn damage.3.2.2 Char Surface SEM After completion of UL-94 testing, SEM images were taken of burnt surface of the samples. The polymer material is insulating so the samples had to be coated prior to SEM analysis. Images of formulation 1 and 3 at 50x magnification are shown below. Theseformulations both have 2.5wt% of MMT and CNF, but formulation 3 has 15wt% FR and formulation1 has no FR. The char formation for these materials is very different because of the intumescent effect of the FR additive. The left side image shows the surface of burnt nylon. The right side image features lighter shaded puffy flakes that appeared when the material experienced combustion. These puffy flakes helped snuff out the ignition of formulation 3.

The right side image features lighter shaded puffy flakes that appeared when the material experienced combustion. These puffy flakes helped snuff out the ignition of formulation 3. The craters visible in the left side image are due to combustion gases building up in the char formation and escaping the material due to pressure differences when it is placed in the vacuum of the SEM instrument.Figure 13: UL-94 char SEM images for formulations 1(left) and 3(right).At the center of the right side image above for formulation 3, there is large agglomerate CNF. This agglomerate is focused in on in Figure 14 at 300x and 1000x magnification. As nylon was burnt away from UL-94 testing, the stronger CNFs were left behind to form these agglomerates. These agglomerates were present in all char surfaces that contained CNF.Figure 14: UL-94 char SEM images of formulation 3 in progressive magnification.4.

These agglomerates were present in all char surfaces that contained CNF.Figure 14: UL-94 char SEM images of formulation 3 in progressive magnification.4. CONCLUSIONSFrom tensile testing, the conclusion is that the lower MMT and CNF percentage, the higher the tensile strength. The differences are not very substantial though. Tensile modulus is higher when the polymer blend has more MMT and CNF. All the blends are very brittle compared to the neat PA11. On the other hand, all blends have higher tensile modulus. Neat PA11 has remarkably high elongation at break. Among the polymer blends, the blends which include 2.5% MMT and 2.5% CNF have higher elongation.In general, mechanical properties of blended nanocomposites with synergistic effects exhibited similar properties to nanocomposites of just one nanomaterial additive.From the SEM image, it is observed that formulation 2, which includes 3.5% CNF, has a higher concentration of CNF than the first one which includes 2.5% CNF.

The MMT clay is too small to be observed by the SEM we used. For the further studies to observe MMT clay, TEM will be needed.From the flammability tests, all the polymer blends were rated UL 94 V-0 except for formulations 1 and 2 which have failed the UL 94 test. Synergistic effect of MMT, CNF, FR additives was observed. Only 15% FR additives is needed with 2.5% MMT and 2.5% CNF to achieve UL 94 V-0 rating as compared to 20% FR additives needed for 5% CNF or 7.5% MMT used individually in PA11 polymer. Flammability properties of these polymer blends need to be further investigated using cone calorimetry.5. REFERENCES1. S. C. Lao, J. H. Koo et al., “Flame-retardant polyamide 11 nanocomposites: further thermal doi: and 10.1177/07349041111404658.2. US Patent Application, 20040021256, dated February 5, 2004.3. S. C. Lao, J. H. Koo et al., “Polyamide 11-carbon nanotubes nanocomposites: Processing, morphological, and property characterization,” Proc.

J. R. Davis, Tensile Testing, 2nd Edition, 1,3, ASM International, 2004.11. C. Gold, Test for Flammability of Plastic Materials for parts in Devices and Appliances, publisher, place, year..

Cells and matrix are selectively assembled inside microchannels using multi-nozzle printing to demonstrate feasibility of chip as a culture environment. The objective of this work is to fabricate a cell-laden microfluidic device by combined solid freeform patterning and replica molding with direct cell writing into channels. This work has application as a 3D physiological model for in vitro pharmacokinetic study in space environment in preparation for long term manned missions.We report on a noveltechnique using additive manufacturing2. IntroductionCell-laden microfluidic devices recreate in vivo cell-to-cell interactions, a current challenge in drug screening, through reproducible quantifiable patterning of spatial and temporal gradients [1-3]. Conventional cell culture of a monolayer on petri dish is a well-established methodology which has been optimized for various drug’s ADMES-Tox profiles in humans [4].

Conventional cell culture of a monolayer on petri dish is a well-established methodology which has been optimized for various drug’s ADMES-Tox profiles in humans [4]. However, the fidelity of conventional models is limited due to local depletion of drug-metabolizing enzymes and transporters, lack of controlled cell seeding and static medium reservoir. Failure to predict drug toxicity by conventional culture methodology is considered the principle reason 90% of new drugs fail during Phase 1 preclinical tests [5]. Biomimetic extra-cellular cues, cell-to-cell contact and definable gradient of drug conditions are engineered into microenvironments for high throughput measurement of toxicity and pharmacokinetics [6, 7].

Cell-laden microfluidic devices improve fidelity of in vitro platforms by physiologically relevant surface area to volume ratio, mass transfer by diffusion, biomimetic microenvironment engineering, dynamic perfusion, and dual organ model [8, 9].The advantage of drug discovery by cell-laden microfluidic devices is reproducible quantifiable control over the stimulation and patterning of biological material [10]. Micron-scale quantities of cell- laden matrix and microfluidic channel produce a cell to interstitial fluid volume ratio of approximately one, as is physiologically relevant [11]. The microfluidic length scale improves uniformity of temperature field and gas supply through cell laden matrix, induce laminar flow pattern and therefore mass transfer by diffusion. These factors reduce variability caused by proportions of experimental design. Hydrodynamic forces on the cells by extracellular fluid produce physical cues stimulate and direct cell function.

These factors reduce variability caused by proportions of experimental design. Hydrodynamic forces on the cells by extracellular fluid produce physical cues stimulate and direct cell function. Dynamic perfusion causes naturally occurring gradients of diffusing chemicals and produces controllable gradient of drug conditions. Cells are assembled in 3-dimensional biomimetic microenvironments to produce biomimetic morphology, cytoskeleton alignment, cell-to-cell contact, focal adhesion of cell to matrix and signaling [12]. Lab-on-a-chip bioanalytic micro-systems introduce multi-component co-cultures to simulate an array of cells by external factors and signals secreted by the cells themselves to engineer a functional tissue.

A controllable gradient of drug captures downstream effects of metabolized drug on target organ and conditions in parallel for physiologically-based pharmacokinetic (PBPK) models [13].Leading microfluidic devices use microfluidic systems and controlled cell seeding to produce cellular microenvironments to facilitate homotypic interaction by co-culture. Vasculature and metatastic prostate cancer microfluidic models use co-culture to understand specific cytokines and growth behavior mechanisms [14, 15]. The effect of stem cells on cancer apopotosis is evaluated in vitro by unidirectional perfusion of a microfluidic device [16]. Microfluidic co-cultures of hepatocytes with nonparenchymal and stem cells are used to study 3-dimensional capillary morphogenesis and drug toxicity [17, 18].

Microfluidic co-cultures of hepatocytes with nonparenchymal and stem cells are used to study 3-dimensional capillary morphogenesis and drug toxicity [17, 18]. In these leading microfluidic devices, microfluidic channels are fabricated in an elastoployer by soft lithography with wet chemical etching or integrated on a polystyrene biochip and cell seeding is controlled by diffusion across matrix, semi-permeable membranes or time course introduction of cell types.Key technology for methodological progress in microfluidic devices for drug discovery include microfabrication techniques and 3-dimensional cell patterning. Miniaturization of perfusion system by microfabrication produces biomimetic interconnected network of channels [19]. Current techniques include photolithography, etching, and molding. Controlled cell assembly is required to produce specific cellular microenvironments and heterogeneous cell-to-cell contact [20].

Current techniques include photolithography, etching, and molding. Controlled cell assembly is required to produce specific cellular microenvironments and heterogeneous cell-to-cell contact [20]. Design models use bottom-up tissue engineering to assemble modular cell environments to achieve biomimetic spatial and temporal interactions of soluble and biological cues. Current techniques include surface patterning by standard photolithography liftoff, photoreactive chemistry, microcontact printing with lamination, molding, and photo-polymerization [21]. Cell migration and separation after printing can be achieved by labeling cells with magnetic nanoparticles or gradient surface treatments to substrate. Cell deposition techniques such as syringe-based cell deposition, rapid prototyping of hydrogel structures, and inkjet-based cell printing include cells as part of the working material to directly assemble biologics in 3-dimensional space [22].

A pumping mechanism is required to precisely transport volumes of liquid to induce specific transient delivery of soluble cues and hydrodynamic strain.A combined solid free from (SFF) patterning and replica molding process offers the following advantages over current techniques to create a cell-laden microfluidic device: All structures and components of a multi-component micro-systems is fabricated in a single step of micro-molding. The channels in the microfluidic network are printed as a 3D continuous pattern. An elastopolymer is cured against the printed pattern and then printed pattern is then cleared from channels thermally or chemically. The process includes direct cell writing to control distribution of the seeded cells in microfluidic device. Cells and matrix are printed directly into channel. The volume and cross-sectional geometry of cell-laden filaments are controlled by manufacturing process parameters.

Cells and matrix are printed directly into channel. The volume and cross-sectional geometry of cell-laden filaments are controlled by manufacturing process parameters. Cell to fluid volume ratio is also controlled by microfabrication and direct cell writing process. The homogenous curing technique of the elastopolymer does not restrict the surface area of the microfluidic device, unlike radiation or point source based exposure techniques. Additive manufacturing techniques increase height of printed mold without also changing width of the mold, allowing for high aspect ratio structures. The printed pattern can be removed from the elastomer microfluidic chip without damaging the channels, which allows for multilevel features and 3D profiles. The design to the finished device takes less than 24 hours to produce.The objective of this work is to fabricate a cell-laden microfluidic device by combined solid freeform patterning and replica molding with direct cell writing into channels.

This work has application for in vitro disease modeling, pharmacokinetic studies and drug screening. Pre-clinical determinations of drug toxicity to liver would greatly improve screening technology [23].3. Method to Produce Cell-laden Channels in Microfluidic DeviceThe combined SFF patterning and replica molding process produces a cell-laden microfluidic device with channel designed to biomimic liver physiology and flow design. First, microfluidic channel pattern is designed from in vivo liver structure and hydrodynamics. Second, microfluidic pattern is printed using precision extrusion deposition. PDMS is cured against the PCL pattern in a single step of replica molding to produce PDMS substrate. PCL pattern is cleared from PDMS, which leaves interconnected microfluidic system. PDMS is plasma treated to decrease surface energy and produce homogenous wetting of the surface during printing. Cell-laden matrix is directly printed into the channels using direct cell writing system.

Cell-laden matrix is directly printed into the channels using direct cell writing system. Cell-laden microfluidic chip is sealed in microfluidic chip and nutrient is perfused through system by programmable syringe pump. Figure 1 diagrams the combined SFF patterning and replica molding for microfabrication of cell-laden microfluidic device process.Fig 1. Combined SFF patterning and replica molding for microfabrication of cell-laden microfluidic device process diagram.Step 1: Microfluidic Channel Pattern to Biomimic Liver Physiology and Flow Design In the liver, diffusion through a 3-dimensional network of microvasculature and bile canaliculi network carries nutrients, waste, and oxygen to parenchymal cells, hepatocytes, and non-parenchymal cells, such as the Kupffer and stellate cells [24]. The relative volume of interstitial fluid to cell matrix is approximately one in liver.

The relative volume of interstitial fluid to cell matrix is approximately one in liver. The microfluidic channel is designed to produce a network of channels to be 50% (v/v) cell-laden matrix and 50% (v/v) interstitial fluid. This ratio is controlled by microfabrication of the channel and direct cell writing process. The range of possible channel dimensions is limited by the manufacturing capability, whose threshold minimum feature size is to be defined by a limiting set of process parameters. The cell volume is controlled by direct cell writing process parameters, and is determined after microfluidic channel geometry is characterized. The microfluidic network is interconnected to allow for dynamic perfusion and controlled transient interaction of interstitial fluid and cell. The system flow rate and architecture cause capillary and hydrodynamic forces acting on the cells.

The system flow rate and architecture cause capillary and hydrodynamic forces acting on the cells. These forces physically strain cell membrane and actin cytoskeleton to induce biomimetic environmental stimulation and cell-to-cell contact necessary to maintain phenotype stability [25]. The microfluidic3.1.channel pattern for 3-dimensional network of microvasculature in liver is simplified to a square wave pattern and elaborated to a sinusoid wave pattern in this work.Step 2: SFF to Produce Pattern SFF by precision extrusion deposition (PED) produces explicitly designed 3-dimensional structures by layer-by-layer deposition of thermoplastic filaments [26-28]. The material delivery system is mounted to a high precision positioning system to print the extruded filament into explicitly designed 3- dimensional patterns.

The material delivery system is mounted to a high precision positioning system to print the extruded filament into explicitly designed 3- dimensional patterns. Design model inputs and manufacturing process parameters, excluding dispensing nozzle diameter, are defined by the operator through automation software Motion Planner (Parker Hannifin Corp). The nozzle is selected from a pre-fabricated set and fixed to the material delivery system prior to extrusion. The material delivery system feeds material into a liquefying chamber by a screw and extruded through interchangeable micron-scale nozzle. Figure 2 presents process information pipeline and system configuration for PED system. Polycaprolactone (PCL) (Sigma) in the form of pellets is loaded into precision extrusion deposition (PED) system’s extruder and heated to 70ºC. Once liquefied and automation software is run, PCL is extruded through 150µm or 350µm nozzle tip by pressure creature by a turning screw at 30rpm.

Once liquefied and automation software is run, PCL is extruded through 150µm or 350µm nozzle tip by pressure creature by a turning screw at 30rpm. Remaining process parameters were derived from previous work using the PED and PCL to produce threshold minimum filament diameter, as it is the limiting case. Nozzle tip is position 0.30mm above level substrate. Single layer square wave pattern of dimensions 20mm x 20mm with struts spaced 1mm on center is produced, as presented in figure 3. Single layer sinusoid wave pattern of dimensions 20mm x 22mm with struts spaced 3mm on center and radius of 1mm is produced, as presented in figure 4. Nozzle is raised 0.30mm between subsequent layers for multi-layer patterns.3.2.Fig 2.

The wettability of cross-linking solution and printed alginate matrix is controllable and homogenous on3.4.post-plasma treated surface. PDMS substrate is placed in RF plasma chamber (Harrick Plasma) for 90 seconds less than 2 minutes prior to printing.Step 5: Direct Cell Writing Into Channel SFF of biological material and matrix by direct cell writing produces explicitly designed 3- dimensional patterns of cell-laden matrix in microfluidic device [32, 33]. The material delivery system is mounted to a high precision positioning system to print the cells into microfluidic channels. Design model inputs and manufacturing process parameters, excluding dispensing capillary tip diameter, are defined by the operator through automation software Motion Planner (Parker Hannifin Corp). The nozzle is selected from a pre-fabricated set of EFD precision capillary tips and fixed to the material delivery system prior to extrusion.

The nozzle is selected from a pre-fabricated set of EFD precision capillary tips and fixed to the material delivery system prior to extrusion. The material delivery system feeds cell-laden matrix through interchangeable micron-scale capillary tip by continuous pneumatic pressure regulated from reservoir of breathing quality compressed air (AirGas). Once pressurized and automation software is run, cell-laden matrix is extruded through 150µm capillary tip 0.5 inches in length by 15psi of air pressure. Capillary tip is position 0.10mm above bottom of channel. Cells are printed in square wave pattern of dimensions 20mm x 20mm with struts spaced 1mm on center to align with PDMS channel pattern. Figure 3 presents system configuration for direct cell writing system aligned to PDMS substrate to print cells into microfluidic channels.3.5.Fig 3.

Cell are maintained at 32ºC and 5% CO2 for less than 3 passages after thawing prior to experiment. Alginate matrix is prepared from 6.0% (w/v) alginic acid sodium salt from brown algae (Sigma) in deionized water and sterilized through aseptic serial filtration through 20µm mesh (VWR). Cells are mixed with alginate for final concentration of 2.0x106 cells/mL HepG2 cells. Prior to printing, channels are flooded with cross-linking solution as formed by 5.0% (w/v) calcium chloride (Sigma) in deionized water.3.7.3.8.Fluorescent Determination of Cell Viabilityis qualitatively evaluated using fluorescence based Live/Dead Viability/Cytotoxicity Kit for mammalian cells (Molecular Probes). Two-color discrimination by the fluorescent probes enthedium calcein AM and ethidium homodimer-1 labels live cells as green and dead cells as red. Samples were analyzed using a DM RIB inverted microscope (Leica) with UV source.

Samples were analyzed using a DM RIB inverted microscope (Leica) with UV source. Images of all cells were captured electronically by Insight 4.0Mp Monochrome digital camera (Spot Insight) and manufacturer provided imaging software.Cell viability post-printing4. ResultsHere we demonstrate the feasibility of a combined SFF and replica molding microfabrication process to produce a cell-laden microfluidic device. Having established control over microfluidic pattern and channel depth, we leverage additive manufacturing techniques to increase aspect ratio of channels and produce 3-dimensional networks of interconnected channels. Results demonstrate process control.Printed PCL Patterns and PDMS Channels Inspection of three designs for fabricated PCL pattern and PDMS channels was performed to demonstrate the pattern could be created, molded, and cleared from PDMS channel.

Figure 4 presents design model input, printed PCL pattern produced by PED system, and PDMS cured against PCL pattern for single layer square wave pattern of dimensions 20mm x 20mm with struts spaced 1mm on center. Figure 5 presents design model input, printed PCL pattern produced by PED system, and PDMS cured against PCL pattern for single layer sinusoid wave pattern of dimensions 20mm x 22mm with struts spaced 3mm on center and radius of 1mm. Figure 6 presents design model input, printed PCL pattern produced by PED system, and PDMS cured against PCL pattern for single layer multi-compartment array of two square and one sinusoid wave arranged in series.4.1.Fig 4. Single layer square wave pattern of dimensions 20mm x 20mm with struts spaced 1mm on center design model input (left) and printed PCL pattern produced by PED system (center) and PDMS cured against PCL pattern with channels flooded with green dye (right).Fig 5.

Single layer sinusoid wave pattern of dimensions 20mm x 22mm with struts spaced 3mm on center and radius of 1mm design model input (left) and printed PCL pattern produced by PED system (center) and PDMS cured against PCL pattern with channels flooded with green dye (right).Fig 6. Single layer multi-compartment array of square and sinusoid wave design models printed PCL pattern produced by PED system (left) and PDMS cured against PCL pattern with channels flooded with green dye (right).Inspection using phase contrast microscope and image analysis software was performed on fabricated samples to characterize PCL pattern and PDMS channel depth and width. Single layer square wave PCL pattern and PDMS chip were viewed from the top and then sliced through the center to inspect cross-section. Figure 7 presents top view and section view of printed PCL pattern and PDMS channels for square wave channel. PDMS curing against the PCL pattern yields a network of channels.Fig 7.

PDMS channel width, depth, and cross-section is controlled by PCL pattern, which is controlled by PCL manufacturing process parameters. The process implements manufacturing control over PDMS channel size.SFF to Incrementally Increase Channel Aspect Ratio Having established PDMS channels can be produced by PCL patterns and the micro-system structures can be controlled by PED manufacturing process parameter, we next investigated the effect of4.2.SFF to increase channel aspect ratio. Inspection using phase contrast microscope and image analysis software was performed on PDMS channel fabricated using 1, 2, 3, and 4 layers to characterize channel width, depth and aspect ratio. A scalpel is used to cut PDMS chip along its centerline to view cross- section of channels. The aspect ratio (AR) is defined as the channel depth divided by the channel width. Figure 8 presents cross section view of PDMS channels molded from multi-layer PCL patterns.Fig 8.

The aspect ratio (AR) is defined as the channel depth divided by the channel width. Figure 8 presents cross section view of PDMS channels molded from multi-layer PCL patterns.Fig 8. Cross section view of PDMS channels molded from multi-layer PCL patterns.The depth of the channel is increased by printing additional layers on top of the first layer. The width of the channel is increased by printing additional layers on top of the first layer. One layer produced a 1.0 aspect ratio channel. Two layers increased the aspect ratio 30%, three layers increased the aspect ratio 60%, and 4 layers increased the aspect ratio 120%.3-D PCL Pattern Cleared from PDMS without Damaging Channel Network Having established PDMS channels can be produced by multi-layer PCL patterns and the micro- system structures can be controlled by SFF manufacturing in 3-dimensions, we next investigated the effect of mutually orthographic layers to create interconnected network of channels.

Inspection using phase contrast microscope and image analysis software was performed on PCL scaffold and PDMS channel fabricated using scaffold to characterize channel width, depth and orientation. Figure 9 presents printed 0-90 PCL scaffold pattern and PDMS channels after PCL cleared from channels.4.3.Fig 9. Printed 0-90 scaffold pattern using two layers as viewed from the top macroscopically and under 4 x magnification using phase contrast microscope (B) from the top (A) and front (B). PDMS Channels after PCL is removed by sonication and the construct is viewed from the top under 4x magnification using phase contrast microscope (C).A porous 3-dimensional PCL scaffold is produced by PED manufacturing. Two mutually orthographic layers are printed by continuous extrusion in the same z-plane. A second set of mutually orthographic layers are printed on top of the first set to produce a scaffold.

Two mutually orthographic layers are printed by continuous extrusion in the same z-plane. A second set of mutually orthographic layers are printed on top of the first set to produce a scaffold. PCL is continuously extrudedduring entire printing episode to produce a continuous pattern. PDMS cures against PCL scaffold to pattern produce a polymer replica with channel network. PCL scaffold is cleared from PDMS chip without disturbing PDMS due to the difference in material melting temperature and solubility. After replica molding PCL is cleared by 50 minutes of sonication in an 80ºC dichloromethane bath. A scalpel is used to cut PDMS chip along its centerline to view cross-section of channel network. Figure 10 presents cross section view of PDMS channels molded from PCL scaffold.Fig 10. Continuous extrusion during PCL patterning creates an interconnected 3-D network for replica molding.Figure 13 presents the top view of PDMS chip and two orthographic cross-section views.

Continuous extrusion during PCL patterning creates an interconnected 3-D network for replica molding.Figure 13 presents the top view of PDMS chip and two orthographic cross-section views. Two layers of channels are produced in PDMS chip. Two rows of pores are visible in both Figure 13 section view A and B. PED manufacturing is used to layer 2-dimensional PCL patterns to create 3-dimensional architecture. This architecture is replica molded and cleared from PDMS chip to produce a 3-dimensional microfluidic system. The depth of the channels from the top of the chip to the bottom of the lower row is 325-385µm.Channels are produced in orthographic directions. The pores visible in the Figure 13 section view A is created by PCL filaments normal to the cutting plane, PCL cleared from the PDMS prior to imaging. Pores in the traverse direction are visible as circular cross-section in Figure 13 section view B.

Process parameters can be varied independent of printed pattern to produce patterns of extruded filaments of variable diameter. The volume dispensed per unit area of substrate is a function of the flow rate and speed of material delivery system. The flow rate through the nozzle is a function of the capillary tip diameter and length, material viscosity, and dispensing pressure. The matrix viscosity is a function of the material and cells concentration, which are selected for to support cell growth after printing. Dispensing tip is selected from a set of prefabricated tips and fixed to material delivery system before printing. Dispensing pressure and speed of material delivery system can be adjusted during printing and tuned to create variable filament diameters. Figure 11 presents matrix extruded through 150µm nozzle tip at 5psi with variable speeds of material delivery system.4.4.Fig11.

Figure 11 presents matrix extruded through 150µm nozzle tip at 5psi with variable speeds of material delivery system.4.4.Fig11. Matrix extruded through 150µm nozzle tip at 5psi with variable speeds of material delivery system.The diameter of the extruded filament, and therefore volume of dispensed material is controllable by manufacturing process parameters. The design input model for direct cell writing system is copied from PED model to produce PCL pattern to trace channel pattern during cell printing. Cell-laden matrix is prepared and loaded into direct cell writing system and printed. Cells are fluorescently labeled after printing to demonstrate their position. Inspection using phase contrast microscope and fluorescent microscopy was performed using image analysis software. Figure 12 presents hepatocytes in alginate printed into channels labeled with the Live/Dead stain and photographed under phase contrast (top row) and fluorescent(bottom row) microscope.Fig 12.

Figure 12 presents hepatocytes in alginate printed into channels labeled with the Live/Dead stain and photographed under phase contrast (top row) and fluorescent(bottom row) microscope.Fig 12. Hepatocytes in alginate printed into channels labeled with the Live/Dead stain and photographed under phase contrast (top row) and fluorescent(bottom row) microscope.Fluorescent imaging confirmed viable hepatocytes were printed into microfluidic channels on PDMS chip. Direct cell writing process produces control over pattern of printed cell-laden matrix and flow rateof extrusion. The process implements manufacturing control over placement and volume of cells dispensed.5. DiscussionThis study describes a novel process to produce microfluidic network based on liver physiology by combined solid freeform patterning and replica molding with direct cell writing into channels.

The process is designed to provide engineering control over environmental cues in microfluidic device; such as microfluidic pattern and channel cross-section, cell to interstitial fluid volume, cell seeding, and perfusion flow rate.Microfluidic pattern is produced by solid free form fabrication of a removable pattern. PED produces continuous PCL patterns of controlled geometry. PDMS curing against PCL pattern produces a polymer replica with channel network. Micro-systems structures of all of the components are formed in a single step of micro-molding. The design model input can be digitally adapted to physiologically relevant patterns and arrays of patterns. Continuous extrusion during the printing process creates an interconnected 3-dimensional network for replica molding. PDMS channels can be produced by multi-layer PCL patterns and the micro-system structures can be controlled by SFF manufacturing in 3-dimensions.

PDMS channels can be produced by multi-layer PCL patterns and the micro-system structures can be controlled by SFF manufacturing in 3-dimensions. SFF techniques during microfabrication process increase channel aspect ratio and embedded architecture in microfluidic system. An interconnected system of mutually orthographic channels was produced by replica molding of a porous scaffold. This capability can be used to introduce hierarchical modular architecture throughout the depth of the chip, in addition to the planar patterning.PDMS channel width, depth, and cross-section is controlled by PCL pattern. PCL pattern is controlled by manufacturing process parameters. Process produces manufacturing control over PDMS channel size and surface area-to-volume ratio oxygen diffusion and thermal field. The diameter of the extruded filament is a function of the nozzle diameter, feed rate (screw speed), temperature, and travel speed of the material delivery system.

The diameter of the extruded filament is a function of the nozzle diameter, feed rate (screw speed), temperature, and travel speed of the material delivery system. In this work, we used two nozzle diameters, 150 and 350µm. Remaining process parameters were derived from previous work using the PED and PCL to produce threshold minimum filament diameter, as it is the limiting case. The filament diameter can be increased by increasing the volume of material dispensed per unit area of substrate. This can be accomplished by increasing feed rate, increasing temperature, or decreasing travel speed of the material delivery system. The process allows for independent control over the filament cross-section and pattern.Cell seeding is controlled by direct cell writing of cell-laden matrix into microfluidic channels. SFF manufacturing by direct cell writing traces channel pattern to dispense cell-laden matrix into channels. The process provides control over cell seeding.

SFF manufacturing by direct cell writing traces channel pattern to dispense cell-laden matrix into channels. The process provides control over cell seeding. The direct cell writing process has multi-nozzle capability to pattern different types of biologics within a single microfluidic device during a single printing episode. The direct cell writing motion system has the capability to move in 3-dimensional space to layer and pattern biologics. Similar to the PED process, direct cell writing can be used to introduce hierarchical patterning of cell/biologics densities or cell types throughout the depth of the chip, in addition to the planar patterning.Direct cell writing process produces manufacturing control of volume of printed material per unit area of substrate and therefore control of cell to interstitial fluid volume ratio.

The diameter of the extruded cell-laden filament is a function of the nozzle diameter, feed rate (dispensing pressure), matrix viscosity, and travel speed of the material delivery system. In this work, we used 150µm capillary tip. Remaining process parameters were derived from previous work printing alginate using direct cell writing process to produce threshold minimum filament diameter, as it is the limiting case. The filament diameter can be increased by increasing the volume of material dispensed per unit area of substrate. This can be accomplished by increasing dispensing rate, decreasing matrix viscosity, or decreasing travel speed of the material delivery system. Since the cross-section of the microfluidic channel has been characterized,operator produce a cell to interstitial fluid ratio close to one by tuning direct cell writing process parameters. The process allows for independent control over the filament cross-section and pattern.

The process allows for independent control over the filament cross-section and pattern. Printed pattern and set of process parameters manufactures a unique 3-dimensional cell-laden construct.Cells are part of the working material in direct cell writing process. The hydrodynamic forces caused by the flowing matrix during printing cause damage to cells. Cell viability in printed construct is a function of the process parameters. During process parameter selection, cell survivability is considered to preserve functionality and viable cell density in printed construct. The speed of the material system is programmable and does not affect hydrodynamic force in the nozzle tips. This means the diameter of the extruded filaments can be controlled without effecting hydrodynamic forces and cell viability. This is desirable, as it produces independent control over filament diameter and cell viability.Perfusion flow rate is controlled by programmable syringe pump.

This is desirable, as it produces independent control over filament diameter and cell viability.Perfusion flow rate is controlled by programmable syringe pump. Precise volumes of fluid in a microfluidic device are transported through microfluidic device through continuous network of channels for medium perfusion, delivery of drug compounds, generation of drug gradient, creation of specific microenvironments and connection of cell culture. Dynamic perfusion using syringe pump and was flow tested to 10µL/hr.6. ConclusionThe objective of this work is to fabricate a cell-laden microfluidic device by combined solid freeform patterning and replica molding with direct cell writing into channels. Process includes SFF manufacturing by precision extrusion deposition (PED) and direct cell writing for independent process control of structural and cell-laden material patterning. Manufacturing process parameters provided control of PCL pattern and cell-laden filament diameter & cross-section.

mbining Additive Manufacturing and Direct Write for Integrated Electronics – A ReviewK. Blake Perez and Christopher B. WilliamsDesign, Research, and Education for Additive Manufacturing Systems LaboratoryDepartment of Mechanical Engineering, Virginia Tech, Blacksburg, VA 24061AbstractDirect write (DW) of conductive materials in the context of Additive Manufacturing (AM) enables embedded electronics within fabricated parts. Previous works use manual, hybrid, and native material patterning systems to deposit conductive materials in parts fabricated by different AM technologies. This capability could eliminate cabled interconnects and redundant electronics packaging, resulting in a significant reduction of mass and assembly complexity. In this paper, the authors explore applications of DW of conductive traces in the context of AM, review prior work in the integration, and analyze the technical roadblocks facing their hybridization.

In this paper, the authors explore applications of DW of conductive traces in the context of AM, review prior work in the integration, and analyze the technical roadblocks facing their hybridization. Barriers to integrating the two technology classes include material, process, and post-process compatibilities.1 INTRODUCTION 1.1 Direct Write OverviewDirect Write (DW) technologies enable the selective deposition and patterning of material. DW is capable of single- and multi-layer, high-resolution, material deposition on both flat and conformal surfaces [1]. DW technologies have been comprehensively reviewed before [2]. Hon and coauthors’ review included methods capable of depositing nonconductive materials. However this work considers only DW methods for conductive materials.

Hon and coauthors’ review included methods capable of depositing nonconductive materials. However this work considers only DW methods for conductive materials. In a review of DW applications, Church and coauthors asserted that DW enables rapid manufacturing of sensors and antennae because it eliminates the masking and etching steps of conventional electronics fabrication [3]. In general, DW processes have been researched and developed because of their potential to replace masked lithographic processes in electronics production.All DW technologies serve the same function (i.e. selective deposition of conductive traces); however, the means of realizing this primary design goal varies dramatically. Extrusion technologies use positive pressure to extrude fluid materials through a small nozzle and onto a substrate. Droplet-based DW ejects small droplets of material onto a substrate.

Extrusion technologies use positive pressure to extrude fluid materials through a small nozzle and onto a substrate. Droplet-based DW ejects small droplets of material onto a substrate. Aerosol Jetting systems aerosolize a material to create a gaseous stream that is aerodynamically focused and deposited on the substrate. Laser-based systems use a laser’s energy to transfer material onto a substrate surface. Finally, tip-based deposition methods use capillary flow of an ink on microscale tip onto the substrate surface [2].When combined with Additive Manufacturing (AM) processes (technologies that fabricate parts by the layer-by-layer joining of material), DW enables the creation of complex and conformal electronics that are structurally integrated into a manufactured part.

When integrated into an AM process flow, DW can be leveraged to manufacture electronic signal routing, embedded sensors, and integrated power systems in additively manufactured structures.1.2 Prior Research and Applications in the Hybridization of Additive Manufacturing andDirect WriteOne of the first examples of a combination of AM and DW is by Palmer and coauthors who produced an electrical interconnects by combining DW and stereolithography(SL) [4]. Other works continued to hybridize AM and DW technologies to fabricate functional electronics on additively manufactures structures.1.2.1 Structurally integrated signal routingBecause of the layer by layer process of AM, electrical interconnects can be structurally integrated into a build. Robinson and coauthors used Ultrasonic Consolidation (UC) and DW to fabricate a cellular aluminum panel with integrated circuitry [5]. The electronics are completely embedded in the robust and lightweight structure.

The electronics are completely embedded in the robust and lightweight structure. Structural interconnects eliminate the need to design for cabling ducts and vias throughout a structure’s interior. The circuit’s encapsulation also protects interconnects from the outside environment.Figure 1 – Cross sectional view of embedded circuit in honeycomb structure fabricated with the hybridization of UC and DW [5]1.2.2 Embedded 3-dimensional antennaeFractal antennas are used to provide a multi or wideband response to different radio signals. Their design consists of fractal patterns to maximize the antennae perimeter while minimizing its footprint. Casanova et al. designed and produced a 3-dimensional fractal antenna using conventional manufacturing and found it to improve radiation efficiency and gain when compared to a 2-dimensional patch antenna [6].

designed and produced a 3-dimensional fractal antenna using conventional manufacturing and found it to improve radiation efficiency and gain when compared to a 2-dimensional patch antenna [6]. Optomec has demonstrated its ability to use the Aerosol Jet DW process deposit on conformal substrates and create functional antennae on AM substrates [7]. One can see the potential of creating 3-dimensional antenna designs with a hybridized DW/AM process.(a)(b)(c)Figure 1 – (a) 3-dimensional fractal antenna [6]; (b) Aerosol Jet of conductive traces on conformal substrate; (c) Aerosol Jet of antenna on AM substrate [7]1.2.3 Conformal ElectronicsConventional fabrication of high density circuitry relies on flat substrates because of process steps such as lithography which require a planar surface. However, products that benefit from complex geometries, such as wearable technology, rely on conformal geometries to better fit the user. Castillo et al.

However, products that benefit from complex geometries, such as wearable technology, rely on conformal geometries to better fit the user. Castillo et al. applied conductive materials to additively manufactured conformal substrates to showcase wearable technology [8]. The helmet insert contains an accelerometer for detecting mechanical shock to the head and can transmit communication of such an event. Figure 2 shows the helmet fabricated using stereolithography, IC components, and manually deposited traces of silver ink [9].Figure 2 – Helmet insert for detecting traumatic head injury [8]1.2.4 BatteriesIn 2008, Malone et al. developed and characterized a Zinc-air battery created with AM technology. The battery was fabricated with a multi-material capable AM extrusion based system [10]. The zinc-air battery is capable of an energy density up to 1.07 mW/cm2. Lewis et al. developed a lithium-ion battery design using extrusion processing with a power density of 2.7mW/cm2 [11].

The zinc-air battery is capable of an energy density up to 1.07 mW/cm2. Lewis et al. developed a lithium-ion battery design using extrusion processing with a power density of 2.7mW/cm2 [11]. The lithium battery is very small (6.6mm3), however it relies on a separatesubstrate with a gold on glass current collector and packaging. Future work could involve the integration with additively manufactured encapsulation.Figure 3 – SEM images of additively manufactured battery electrode [11]1.2.5 Discreet ElectronicsAM and conductive materials can also be utilized to discreetly implement electronics into common or unsuspecting items. This application enables discreet and unobtrusive data collection. Medina et al. used stereolithography and DW to place a small camera and video transmitter within the internal structure of a printed alarm clock [12].

Medina et al. used stereolithography and DW to place a small camera and video transmitter within the internal structure of a printed alarm clock [12]. The camera embedded in the alarm clock showcases the ability to leverage AM’s geometric reproduction abilities and the utility of conductive traces to create discreetly embedded electronics.(a)(b)(c)Figure 4 – (a) original alarm clock; (b) alarm clock internals; (c) printed base to incorporate wireless camera module1.3 ContextWhile technologies and applications of both AM [13] and DW [14] have been independently reviewed before, the literature analysis presented in this paper is specifically focused at the intersection of AM and DW. Previously, Lopes and coauthors identified the opportunity for expanding AM to include electronics systems [15]. Their implementation is narrow in the context of utilizing only Stereolithography and extrusion-based deposition.

Their implementation is narrow in the context of utilizing only Stereolithography and extrusion-based deposition. In this work, the authors analyze different combinations of DW and AM processes to evaluate the tradeoffs considered in selecting the technologies and materials as well as their resulting performance characteristics (e.g.,feature size, conductivities, and material-substrate compatibilities). Section 2 details the DW processes most applicable to hybridization with AM technologies. Section 3 addresses roadblocks and challenges that are present in integrating AM and DW processes. Closure is offered in Section 4.2 DW PROCESSES FOR AMIn order for DW technologies to integrate with AM processes, it is paramount that the material deposition system is maneuverable and reliably positioned. Additionally, for manufacturing relevance, the write speeds should be comparable to the corresponding AM process.

Additionally, for manufacturing relevance, the write speeds should be comparable to the corresponding AM process. Of these technologies reviewed by Hon et al., Aerosol Jet, inkjet, and extrusion methods are considered immediately relevant for hybridization with AM because they afford freeform control of both the deposition tool and the material delivery system. Basic performance capabilities of these DW technologies are summarized in Table 1 and summarized in the following sub-sections.The remaining DW technologies were excluded from this review due to their slow write speeds, incompatible reaction environments, and large deposition tool heads or material precursors that are difficult to integrate with the AM environment. As both DW and AM technologies evolve, excluded DW processes may become relevant in the context of hybridization.

As both DW and AM technologies evolve, excluded DW processes may become relevant in the context of hybridization. These other technologies are sufficiently explained and analyzed in previous reviews of DW technologies [2], [14].Table 1 – Comparison of direct write technology capabilities. Adapted from [2]TechnologyMinimum ResolutionViscosity RangeMax Write SpeedInkjet≥20 μm≤0.1 Pa∙s0.30 mm3/sAerosol Jet≥10 μmNot Applicable0.25 mm3/sExtrusion≥25 μm≤5000 Pa∙s300 mm/s2.1 Extrusion-based DepositionExtrusion-based DW deposition uses positive pressure to dispense conductive materials in fluid form through a small nozzle. Both syringe (Figure 5a) and micro-dispensing pump (Figure 5b) systems fall in this category. The pneumatic control of these systems allows for dynamic operation and pressure regulation.

Both syringe (Figure 5a) and micro-dispensing pump (Figure 5b) systems fall in this category. The pneumatic control of these systems allows for dynamic operation and pressure regulation. Maintaining precise control over pressure is important as pressure directly affects feature width, even if ink type and nozzle diameter are held constant. These systems can also apply a vacuum to draw fluid into the dispense tip and mitigate dripping or over dispensing. The distance from which material can be deposited from the substrate can vary; however, it is understood that for different machine parameters (i.e. tool speed, material type, pressure), there is an optimal offset for the dispense tip [16], [17].(a)(b)Figure 5 – (a) schematic of basic syringe system (b) micro-dispensing pump [16]Extrusions systems can dispense fluidized materials with viscosities up to 5000 Pa∙s with feature widths as low as 25 μm.

Linear writing speeds can reach 300 mm/s, although 50 mm/s is typically observed [2]. Volumetric rates depend on nozzle size, which is limited by the material selection.These systems are most typically used to dispense metal-based inks. Metal-based inks consist of metal particles or flakes dispersed in a volatile solvent that evaporates after being dispensed. The metals most commonly used are silver and gold because of their resistance to oxidation; however, this can be cost prohibitive (e.g. 70% silver by weight inks can cost about $2 USD per gram). Other formulations use metal particle suspended in a semi-conductive matrix as used in [18].Carbon based inks can also be dispensed with these systems although they are less conductive than metal inks. Leigh et al. showed a measured a conductivity of 103 S/m, two orders of magnitude less conductive than unsintered metals [19].

Leigh et al. showed a measured a conductivity of 103 S/m, two orders of magnitude less conductive than unsintered metals [19]. Carbon based conductors could find an application in an education context where cost is more critical than electrical performance. Carbon based conductive inks cost less than $0.50 USD per gram.For DW purposes, the metal content of the ink is a critical factor. Higher loading results in higher particle density as deposited and thus better conductivity. However, inks with high metal loadings are more viscous and difficult to dispense. Viscous inks require more pressure to dispense and their high particle content leads to clogging issues in the dispensing orifice. Clogging is alleviated by using larger nozzle tip diameters, although this negatively impacts the minimum achievable feature size of the system. Most importantly, high metal loadings improve conductivity but limit the minimum achievable feature size.

Most importantly, high metal loadings improve conductivity but limit the minimum achievable feature size. As deposited, these metal-based inks can achieve a conductivity in the range of 105 S/m [17]. If these materials can be exposed to highertemperatures to undergo sintering, the conductivity approaches 107 S/m, the conductivity of bulk metals. This temperature processing is discussed in depth in section 3.1. Ideal inks balance loading for clog-free dispensing and desired feature resolution while maximizing conductivity.Extrusion processes are favored for dispensing heavily loaded inks as they are capable of extruding such viscous material. Previous works have used inks with metal loadings in the 60-70% range. From a control perspective, extrusion systems are favorable for integration with AM technologies because the tool tip is maneuverable, can dispense in different orientations, and can process high volumes of material.

Additionally, the 50 μm feature resolution is comparable to contemporary AM technologies, although this value increases with the use of particle loaded inks.2.2 Ink JettingSimilar to technology in inkjet printers, Inkjet DW is a droplet-based technology that places material where it is needed. Some Inkjet DW systems use a piezoelectric diaphragm and a small orifice to dispense material in liquid form. As the diaphragm expands, it places positive pressure on the fluid, which is expelled through the orifice as shown in Figure 6. Alternatively, some systems use a heat source to heat the ink and generate a bubble. This bubble provides the pressure to expel the droplet through the orifice.

Alternatively, some systems use a heat source to heat the ink and generate a bubble. This bubble provides the pressure to expel the droplet through the orifice. Thermal systems however can cause solvents in an ink to volatize rapidly, leaving behind a high concentration of particles which can clog the printing orifice.Figure 6 – Schematic of ink jet deposition [13], [20]Inkjet systems can dispense fluid materials up to 100 mPa and can create features as small as 20 μm. The volumetric dispense rate of a single nozzle is only 0.3 mm3/s, however this can be increased by using an array of nozzles. The formation of droplets and their corresponding size is heavily influenced by the distance of a print head from the substrate [14]. Deviations from an optimal distance can result in poor resolution control.

Deviations from an optimal distance can result in poor resolution control. Unlike extrusion systems, inkjet technology has drop by drop control allowing for more discrete placement of material.Like extrusion technologies, inkjet systems use metal or carbon particles suspended in a volatile solvent; however, the inks must be less viscous to be compatible (i.e. lower solids loading) [14]. As with the extrusion-based systems, the particle loading plays a significant role in theimply higher dispense characteristics as well as conductivity. Higher concentrations conductivities; however, high concentrations make the ink more viscous and therefore difficult to jet. When compared to extrusion-based systems, for the same orifice diameter, inkjet systems are more sensitive to particle loading and clog at lower concentrations. A thorough review of droplet formation and viscosity limitations for ink jetting is found in [21].

A thorough review of droplet formation and viscosity limitations for ink jetting is found in [21]. Typical formulations contain 30-40% metal by weight although up to 50% loadings can be found. These higher percent loadings implement very small particle sizes which faces the challenge of agglomeration and reduces jetting reliability[22].Inkjet systems are also well suited for dispensing metallo-organic decomposition (MOD) inks. MOD inks contain no particles. Instead the metals are dissolved in solvent such as toluene or xylene [23]. When exposed to air, the solution decomposes leaving the metal precipitate. This reaction time can be longer than the dry times of solvent based inks. MOD inks’ viscosities can be tailored specifically for ink jetting—favorable for high resolution capabilities [24]. The metal content of these inks is typically only 20-25% by weight. Previous works have used multiple printing passes order to deposit more metal and improve conductivity.

The metal content of these inks is typically only 20-25% by weight. Previous works have used multiple printing passes order to deposit more metal and improve conductivity. As with any metal ink, post- process sintering can dramatically improve conductivity. Both metal-loaded and MOD inks are compatible with ink jet technology. Choosing a solution depends on the importance of metal content (higher with metal-loaded inks) versus the rheological customization (and thus resolution) of MOD inks.Ink jet deposition heads are maneuverable and can also print in different orientations making it a good candidate for AM hybridization. Ink jetting can be tailored to produce traces with high resolution in arrays that are comparable to AM processes in terms of throughput. Compared to extrusion systems, ink jetting is capable of higher resolution. However, lower achievable metal loadings negatively impact conductivity.

Compared to extrusion systems, ink jetting is capable of higher resolution. However, lower achievable metal loadings negatively impact conductivity. Additionally, ink jetting is currently used to deposit inks for pulsed photonic sintering which may be a sintering solution that is compatible with low temperature substrates. This technology is discussed further in section 3.1.2.3 Aerosol JettingAerosol Jet DW systems aerosolize solutions into a vapor stream to deposit material. For the DW of conductive materials, the same metal-solvent inks can be used. This mixture is dispensed through a high velocity nozzle as shown in Figure 7. The vapor accelerates through the nozzle and impacts the substrate surface. The aersolization process can help to dry out the solvent constituents of the solution and decrease dry times. The size distribution of the particles can be controlled independently of the feedstock’s distribution using virtual impactors.

The size distribution of the particles can be controlled independently of the feedstock’s distribution using virtual impactors. A virtual impactor uses a secondary gas flow with a different flow rate. Smaller particles are diverted by the secondary flow. Larger particles’ momentum prevents them from being diverted and continue on the main path. This allows for a homogenous mixture of uniformly sized particles, which enables high resolution and consistent post-processing outcomes (discussed further in Section 3.1).Figure 7 – Optomec Aerosol Jet process [25]Aerosol Jet DW is capable of producing features as small as 10 μm with linear write speeds of 300 mm/s although 50 mm/s is typically observed. Single passes can be as thick at 2 μm depending on the ink [7]. Therefore at a resolution comparable to extrusion system (50 μm), the volumetric deposition rate is limited to 0.03mm3/s.

Single passes can be as thick at 2 μm depending on the ink [7]. Therefore at a resolution comparable to extrusion system (50 μm), the volumetric deposition rate is limited to 0.03mm3/s. Because of the aersolization process, viscosity is not so much an issue as long as the material can be aerosolized (includes metal and carbon). The deposition head can also deposit material at a range of distances from the substrate surface. Deposition is possible with a standoff distance within the range of 1mm to 5mm [14]. This makes motion control for deposition on conformal surfaces simpler.Again the maneuverability of the deposition head and the systems control make Aerosol Jetting a good candidate for integration with AM systems. The tolerance of the dispensing nozzle’s standoff distance makes the system flexible and ideal for deposition on conformal surfaces.

The tolerance of the dispensing nozzle’s standoff distance makes the system flexible and ideal for deposition on conformal surfaces. Flat surfaces or high volume production requirements however are better suited by the other technologies given the Aerosol Jet deposition rate is relatively low.3 ROADBLOCKS AND SOLUTIONS THROUGH ANALYSIS OF PREVIOUS WORKSThe hybridization of DW and AM is non-trivial because of (i) material post-processing temperature compatibilities, (ii) adhesion between conductive and substrate, (iii) adhesion between conductive layers, and (iv) adhesion to external circuitry. Potential solutions are presented in the context of an analysis of previous works encountering these roadblocks.3.1 Temperature ProcessingMost of the work done in the past decade in combining AM and DW has been confined to selecting AM materials that can withstand high-temperatures and DW materials that require relatively low post-processing temperatures.

For example, most metal DW inks require post processing at a minimum of 100°C to become conductive. Only a few polymeric AM materials can withstand these temperatures. Some of these materials include Duraform HST (SLS),Prototherm (Stereolithography), and ULTEM (FFF). Other AM processes like PolyJet have materials that can only withstand a maximum of 80°C.High temperatures allow individual particles to form necks and eventually grain boundaries between particles. This neck and subsequent grain boundary formation increases contact area and, consequently, conductivity. The temperature at which diffusion and sintering occur is different for different metals. This is typically at least 70% of the material’s melting temperature which is near 1000°C for metals. The melting temperatures of metals in particle decreases when the particles are sufficiently small [26]. Figure 8 demonstrates how melting temperature decreases as particle diameter also decreases.

The melting temperatures of metals in particle decreases when the particles are sufficiently small [26]. Figure 8 demonstrates how melting temperature decreases as particle diameter also decreases. One conclusion is that sufficiently small particles on the nanoscale could sinter at low temperatures compatible with AM substrates.Figure 8 – Melting temperatures as a function of particle radius demonstrating lowered melting temperatures as size decreases [26]As the diameters decrease into the nanoscale, inter-particle forces begin dominate the suspension of the particle-solvent mixture. Maintaining a homogenously dispersed mixture becomes a challenge. Resulting agglomerates of the particles are larger than their singular, dispersible, counterparts. Agglomerates are difficult to dispense without clogging. It is possible to coat the metal particles with capping agents to prevent agglomeration; however, the capping agents themselves require high temperature processing for removal.

It is possible to coat the metal particles with capping agents to prevent agglomeration; however, the capping agents themselves require high temperature processing for removal. The capping agents then become the limiting factor for post-processing temperature.Figure 9a shows particles still encapsulated by the capping agent that prevents agglomerationswhile in solution form. At this stage, the materials is only slightly conductive, if at all.b exhibits the interfacing of the particles after the capping agent is removed. Temperatures ofat least 100°C are typically required for this.c demonstrates low temperature sintering. Based on Ostwald ripening, surface particles have migrated to more stable sites, lowering the overall surface energy [27]. This results in densification and more contact area between particles, however at its limit the resulting structure is still porous and thus not as conductive as the bulk material.d shows a fully sintered and dense formation.

This formation reaches the conductivity of the bulk metal however it requires temperatures of 70-90% of the materials melting temperature. Perelaer et al. provide a more extensive review of conductive inks and the factors governing their post-process requirements [28].(a)(b)(c)(d)Figure 9 – Different stages of post-processing to improve conductivity of particle-based metal inksMultiple works by Medina, Wicker, & Lopes have integrated Stereolithography and DW technologies to create an array of functional electronics printed onto additively manufactured structures. All of these examples take advantage of high-temperature SLA resins that can withstand the 110°C post-processing temperatures required by their inks [5], [8], [9], [15], [17], [29].While not focused on AM, there has been interest in evolving sintering techniques to be compatible with low temperature substrates, specifically polymers.

Perelaer and Schubert have investigated multiple localized methods for sintering directly written conductive traces [28]. They found that material can be locally sintered using laser or microwave radiation. Laser based sintering still poses the risk of damaging a polymer substrate. This damage on an SLA substrate is characterized by Medina et al [17]. Microwave radiation based sintering exhibited promising results; however, microwave conductors or antenna were required to absorb useful amounts of radiation which is unpractical for volume processing. Another method for polymer compatible sintering is pulsed photonic curing. Pulsed photonic curing is interesting in the context of AM because it can heat metal traces to high temperatures without damaging the surrounding substrate. It is also a broadcast energy source so alignment to the traces is not necessary [30].

It is also a broadcast energy source so alignment to the traces is not necessary [30]. Preliminary work by Cormier suggests that the technology is safe for low temperature AM substrates [31].(a)(b)Figure 10 – Photonic curing of copper on (a) FFF and (b) SLS substrates. (b) Also shows an intentionally uncured region in black [31]Figure 10 shows examples of copper cured by photonic curing on both FFF and SLS manufactured substrates. The conductive materials for photonic curing can easily be deposited using inkjet technology and the cured material and results have shown conductivities of 109 S/m which is greater than bulk conductivity of copper. The disadvantages of photonic curing are that the distance from the energy source is critical, therefore curing conductive inks is restricted to planar surfaces. Additionally, the thickness of the deposited ink is critical for curing and is restricted to relatively thin layer heights.

Additionally, the thickness of the deposited ink is critical for curing and is restricted to relatively thin layer heights. Successful curing has not been demonstrated with thicknesses larger than 1-3 microns [30].3.2 Adhesion of DW materials to AM substratesFor any deposition process, adhesion between the deposited material and substrate is important. Poor adhesion between the materials can result in delamination of the conductive materials from the substrate and failure. This is especially important for traces that are embedded within a print and not on the surface. After a circuit is embedded, access and repair is not possible. It is difficult to discern without testing if the adhesion between particular inks and AM substrates is possible.

After a circuit is embedded, access and repair is not possible. It is difficult to discern without testing if the adhesion between particular inks and AM substrates is possible. Factors influencing adhesions include:Surface energy of both DW material and AM substrate  AM substrate surface roughness  AM substrate surface treatment  Chemical interaction between DW & AM materialsThe surface energy of the DW material and AM substrate can determine the contact angle between the materials or surface wetting. Better wetting implies more contact area, more surface energy and thus better adhesion. The surface roughness of the AM substrate also can affect the adhesions characteristics. Cheng et al. determined that the surface roughness could be especially important to the adhesion of microparticles on a substrate, A more detailed explanation of thisinteraction can be found in [32].

determined that the surface roughness could be especially important to the adhesion of microparticles on a substrate, A more detailed explanation of thisinteraction can be found in [32]. Surface treatments such as etching, plasma treatment, and priming can change the surface tension of substrates to enhance adhesion characteristics [33], [34]. Finally, a chemical interaction between the DW material and AM substrate can also contribute to the adhesions characteristics of the parts. Understanding these parameters is important to achieve good adhesion.Some AM technologies pose unique challenges with respect to surface adhesion because of the surfaces that they produce. Technologies like fused filament fabrication (FFF) have a very non- uniform surface on the macro scale which poses a unique problem for DW.

Technologies like fused filament fabrication (FFF) have a very non- uniform surface on the macro scale which poses a unique problem for DW. FFF produces a ridged surface because its material patterning process.Figure 11 – Close up image of silver inks deposited on FFF ULETM 9085 printed substrate [35]Figure 11 is a magnified image of silver ink deposited on an FFF surface. Since the surface is not flat, conductive inks tend to fill into the valleys between the AM traces resulting in short circuits or discontinuities. Optomec and Stratasys utilized a UV curable dielectric material as an intermediate bonding layer between the FFF part surface and the conductive ink [35]. The dielectric material fills the voids and dries with a planar surface as shown in Figure 12.Figure 12 – Conductive ink deposited on FFF substrate using dielectric intermediate bonding layerThe dielectric material creates a flat surface for depositing conductive material.

The dielectric can also be chosen with good adhesion characteristics to both the substrate and conductive materials. For this reason, an intermediate layer can also be used with other AM processes such as Stereolithography to benefit adhesion even if the surface is much smoother than that of the FFF process. Because of the flexibility of DW systems it is possible to use the same systems for depositing the dielectric material.Moreover, an intermediate bonding layer can help to alleviate stresses induced by a mismatch in the CTEs of the AM and conductive materials which is common between a metal and a plastic. Sintering elevates the temperature of the conductive material and cracking may occur if the difference in CTEs is sufficiently high as discussed in section 3.1.3.3 Interlayer bonding and conductivityInterlayer bonding is an important aspect of DW to consider in the context of AM because a multilayered process is inherently more compatible with AM process.

Moreover, it enables more complex circuit geometries and configurations within a part. Previous research regarding DW of circuits has largely focused on planar, single layer, deposition because of its applications in PCB fabrication. Many of the hybridizations of DW and AM so far have only incorporated these single layer DW methods. While DW has been shown feasible on 3-dimensional conformal surfaces, the deposition is still done in a single layer [9].Lopes and coauthors have shown DW compatibility in three-dimensions with deposited vias [15]. This process combines SLA and extrusion-based DW. Figure 13 diagrams the process. First, IC components are embedded in a partially completed build. A second build process creates vertical interconnects for connecting the embedded components. The stereolithography build finishes with traces to connect all the embedded components. The extrusion-based DW process finally dispenses conductive ink in the traces and into the vias.

The stereolithography build finishes with traces to connect all the embedded components. The extrusion-based DW process finally dispenses conductive ink in the traces and into the vias. The vertical interconnects are explicitly cited as a potential manufacturing and design issue. They require direct access from the surface and complex channels can be difficult to fill with ink. In situ multilayer deposition would allow for conductive materials to be deposited layer by layer in conjunction with the corresponding AM process and alleviate the design complications of the vias with designs that do not require access from the surface. However bonding between the layers of conductive inks is a critical issue.Figure 13 – Build stages for DW and SLA incorporating vertical interconnects.

However bonding between the layers of conductive inks is a critical issue.Figure 13 – Build stages for DW and SLA incorporating vertical interconnects. (a) sockets for embedded components; (b) vias to access embedded components (c) planar interconnects [15]Multilayer DW is desired in the hybridization of DW and AM to leverage the full capabilities of the AM process. Layered DW allows for multiple layers of circuitry in a single build as well as conductive features orthogonal to the build plane. However, it has been shown that in layer based deposition of conductive materials, the Conductivity along the plane of depositions is higher than conductivity between layers, perpendicular to the plane of deposition. [19]. This is due to lower contact density between layers versus along layers. Good interlayer bonding minimizes this.A potential solution to promoting interlayer bonding and hence conductivity is a graded volatility solvent. Sun et al.

Good interlayer bonding minimizes this.A potential solution to promoting interlayer bonding and hence conductivity is a graded volatility solvent. Sun et al. have shown this type of solvent to promote adhesion between layers so that printed structures can withstand sintering without delamination and distortion [11]. Water is used to enable the deposited layer to dry enough to solidify and maintain shape. Ethylene glycol and glycerol are also part of the solvent to maintain moisture after the water has evaporated and promote interlayer bonding. Once the structure is finished, the other solvent constituents can be removed with temperature processing. This is one potential solution to creating solvent based inks that are compatible with AM processes yet maintain conductivity.4 CONCLUSIONSThis paper evaluates different DW technologies specifically well-suited for hybridization with AM technologies.

The hybridization of DW for conductive materials and AM creates a new class of manufacturable technology that integrates electronic signals and sensing within the structure of a manufactured part and leverages the design freedoms enabled by AM. The applications and benefits of this have only partially been exposed but are plenty and promising. As both technologies are additive processes, there is a significant reduction in waste when compare to traditional machining and electronics fabrication processes. Moreover, the versatility of both classes of technologies allow for a significant reduction in process chain complexity through a reduction in processing steps. These hybridizations have application in both prototyping and high- volume production spaces.

These hybridizations have application in both prototyping and high- volume production spaces. Understanding the metrics that make certain AM and DW processes compatible is paramount in developing successful pairings of the two technologies.The previous works discussed in this paper have provided a foundation to understanding the capabilities and limitations of combining DW and AM systems. While the functional aspects are well demonstrated, relevant metrics for the electrical performance of these devices are not fully evident. More process modeling and optimization remain to be done in order to understand the adhesion between DW and AM materials as well as analysis of achievable conductivities of different materials on various substrates.While this paper focuses on the hybridization of two distinct technologies, it is prudent to understand how these technologies could come together as a unified solution.

An example of this is the carbon-filled FFF material developed by Leigh and coauthors [19]. Although not as conductive as the metal materials used in other works, the carbon material is processed through the AM technology’s native material patterning system allowing the deposited conductive material to match the parameters of the structural components. An AM process that is capable of depositing both structural and conductive materials is ideal. But by matching the parameters of both DW and AM systems, it is feasible to create a hybridized system that behaves more as a unified technology. Compatible process resolution, throughput, and temperature capabilities in conjunction withconductive and substrate materials matched for quality adhesion defines the most desirable case of hybridization.5 ACKNOWLEDGMENTSThe authors thank the Air Force Research Laboratory for their financial support. Additionally,I thank Dr.

MBINING ADDITIVE MANUFACTURING WITH COMPUTER AIDED CONSUMER DESIGNY. Ariadi, R.I. Campbell, M.A. Evans and I.J. GrahamDesign Practice Research Group, Loughborough Design School, Loughborough University, Leicestershire LE11 3TU, UKAbstractThis paper reports an investigation into the potential for consumers designing and manufacturing their own products using a combination of “Computer Aided Consumer Design” (CaCODE) and Additive Manufacturing (AM). Recent developments in the field of AM (cheaper machines and new materials) have led to renewed interest in the manufacture of customised products and, more specifically, allowing consumers to create their own bespoke products. However, a persistent weak link in this paradigm is the inability of most consumers to create 3D models as an input for AM. Operating a conventional CAD system requires a lengthy period of specialist training and is therefore not viable in this context.

Operating a conventional CAD system requires a lengthy period of specialist training and is therefore not viable in this context. Consequently, easy-to-use 3D design tools are needed to make AM more accessible to consumers. This research study investigated the suitability of such a system for enabling consumers to design their own pens for manufacture using AM. The investigation also explored the consumer acceptance of current AM capabilities when used for the production of consumer products. The results showed that careful attention must be paid to the specific needs of consumers, both in terms of their product preferences and their ability to use software. These will be used to guide the design of future CaCODE systems.IntroductionThe overall research aim of this project is to investigate the potential for consumers to actively engage in product design and manufacture.

This is done regularly for aspects such as options configuration or choosing colours, but not so often for 3D shape generation. A key reason for this is the inability for ordinary consumers to create or modify 3D models. To provide an appropriate system for consumers to produce 3D models for AM, a so-called Computer Aided Consumer Design (CaCODE) approach has been employed. Consumer design refers to products whose conception, specification, design, or manufacture may occur with direct consumer input [1]. The driving philosophy of CaCODE is that it must be relatively easy to use, as it will have to be operated by the members of the general public who do not have any CAD skills.Since the role of the CaCODE software is to generate 3D models that will ultimately be used to create actual products, this study also reviewed the current capabilities of various AM technologies when producing a consumer product (using an asymmetric pen as a case study).

The pen was selected as an example of a consumer product because of its simplicity and the fact that consumers can readily identify with it. The acceptability of AM products to consumers was measured in terms of surface finish, colour quality, weight and grip.238BackgroundDesign activity has been described as everything from “the teacher arranging desks for a discussion to the entrepreneur planning a business to the team building a rocket” [2]. In the areas of Mass Customisation (MC), personalisation or individualisation, the terms “user co-design” or “customer integration” are familiar [3]. There is also an opinion that “with user-design systems, the professional designer is replaced by the user,” [4]. Research at Loughborough Design School has indicated that such wholesale replacement is not currently feasible and that collaborative consumer design is required, where part of the product design is done by the designer and the remainder by the consumer.

As part of this approach, this paper will evaluate potential software and hardware systems to see how suitable they are for bringing consumers into the design process and then progress through to manufacture. The following sections review the direction of designing and manufacturing that is usually done by companies but which might, in the future, be undertaken by consumers.The Increasing Role of the ConsumerMany industries have shifted from production-driven and market-driven approaches to consumer-driven approaches. This latter approach connects consumers’ choices with the capabilities of the company and extends the philosophy of concurrent engineering to sales, marketing and end users. Thereby, it brings the voice of customers into design and manufacturing [5].The widespread use of the personal computer and its peripherals, as well as the Internet and e-business, has affected the behaviour of many consumers [6].

Communication technologies allow easy and fast interaction between manufacturers and consumers so manufacturers can rapidly identify consumer requirements and consumers can express their particular needs [7]. In this way, consumers are becoming accustomed to manufacturers, so that the numbers of consumers who want to have personalised items has the potential to increase significantly.to conveyingtheir requirementsBringing Back “Craft Customisation”In the pre-industrial era, craft customisation existed. People created products in a customised way despite the limitations of technology at the very beginning of market trade. Products were initially made one-by-one according to each individual’s needs. This procedure was typically carried out by individuals or home industries. The desire to reduce production costs and time created strong influences in the increased use of Mass Production (MP). Good quality and a cheap price became very popular drivers.

The desire to reduce production costs and time created strong influences in the increased use of Mass Production (MP). Good quality and a cheap price became very popular drivers. MP provided a low cost option that led to uniformity. This drove craft production into the realms of art and bespoke products, where higher cost and product variability were not such important issues. At some point, simplification of product variants was criticised, and consumers required industries to listen to their expectations of styles, sizes, needs or even schedules. This drives production away from pure MP towards the need to accommodate versions and options. Going beyond this, MC is a production system that enables customisation and personalisation or individualisation of products as well as services at a price comparable to MP [8].239AM technology can move the method of customisation away from the conventional means, such as product modularity, towards bespoke production.

AM capabilities can serve consumers as individuals in the same way craft customisation can, but employing different forms of communication and interaction. In addition, with this relatively simple type of manufacturing (no tools, little set-up) the manufacturing can be undertaken by SMEs or even individuals. In some ways, this shows similarity to the manufacturing conditions of the pre-industrialisation era. According to Davis, similar events to those in the past will happen in the future regardless of how or what form they will take [9]. Hence, the use of AM could be portrayed as a means of completing the circular pattern of production technologies shown in Figure 1.Figure 1: Production Technology cycle: from craft customisation to AM customisationAlongside the movement of AM into smaller organisations, CAD has already progressed in a similar direction.

Historically, CAD was first developed for large scale manufacturing companies, such as those found in the automotive and aerospace sectors. Early computer aided design tools were not easy to operate, even for well-trained design engineers. With further developments, CAD has taken many steps to become an industrial design tool also. With regards to opening up AM advantages to more people and to bringing manufacturing closer to consumers, this research focuses on extending the use of CAD to consumers. If feasible, then consumers themselves could modify digital models without the need for extensive CAD experience. Consequently, this would place a new design tool on the market for consumers (Figure 2).Figure 2: The progression of CAD towards the consumer240Consumer innovation, personalisation and sustainabilityFor four decades, several studies have observed consumers’ innovative buying behaviour.

These studies focused on consumption values to explain why consumers choose to buy or not to buy certain products, as well as the relations between consumers and products. As a result, four types of motivation underlying consumer innovativeness were identified, namely functional, hedonic, social and cognitive motivation as shown in Table 1 [10].

The anticipated advantages of CaCODE combined with AM are to enhance the motivation for consumer innovation.Functional Usefulness, handiness, compatibility, efficiency, comfort, ease, quality, reliability.Hedonic Pleasure, fun, sensation, excitement, enjoyment, tension, desire, an escape from the daily round.Social Being different and unique, status, standing, prestige, distinction, opinion leadership, manipulation, visibility, social rewards, trendiness, symbolism, demonstrating one's success, sense of belonging, image.Cognitive Knowledge, information, intelligence, wisdom, eagerness to learn, logical thinking, insight and understanding, reason, brainpower, mental stimulation.Table 1: Examples of four consumer innovation motivationsProduct personalisation refers to a process that defines or changes the appearance or functionality of a product to increase its personal relevance to an individual [11].

Consumers need to invest mental effort in a product to configure choices of shapes and colours. As a result, the effect of the effort invested during the customisation process generates an emotional bond with the product. It has been shown that users will extend the period of time spent with the product [12]. This can postpone people's tendency to replace products for non-technical reasons and the longer consumers keep products the better it is for the environment, in terms of product disposal or the energy going into recycling. In this way, it is anticipated that greater consumer involvement in design will lead to improved product sustainability.Research DesignThis research study employs quantitative and qualitative methods integrated through participant surveys, in which the participants will be asked to evaluate the use of a CaCODE tool and assess the appearance of AM products.

Both types of data will likely provide useful information such as user acceptance of the method and products. It will also help to narrow down possible directions for follow-up research as well as improvements to future CaCODE systems. The authors developed a CaCODE pen design tool for this survey and built pen samples using several AM systems, coupled with different finishing techniques. Non-designers from a variety of backgrounds were asked to design their own pens using the tool and then to evaluate the AM products as if they were the results of their design input. Consumers were then asked to evaluate the software's user-friendliness and their acceptance of the sample products.241Developing CaCODE:PenOne type of consumer design tool available today is design configuration programs. These programs are also called co-design tools and defined as software applications that create particular products where consumers can build a certain version by choosing some of its elements.

These programs are also called co-design tools and defined as software applications that create particular products where consumers can build a certain version by choosing some of its elements. This type of tool is limited by the choices of design configuration provided and can often be confined to choosing colours, patterns and performance options.Recently, web-based “easy CAD” tools have been released, which consumers can access through the Internet with no installation required. These are 3D CAD systems and users are required to build 3D models from scratch. A certain degree of design skill is still required, which can be a problem, since some people believe that not everyone can be a designer [13], [14]. To overcome this problem, this research offers an easy CAD system that does not require consumers to start from a blank screen. Instead, they are given a starting point product that they can manipulate until they reach shape with which they are happy.

Instead, they are given a starting point product that they can manipulate until they reach shape with which they are happy. Therefore, it will fill the gap between the existing design tools (design configuration tools and blank screen CAD) for consumers in terms of producing 3D data for printing AM products (Figure 3).Figure 3: Filling the gap between design configuration tools and blank screen CADTo simplify the operation of CaCODE, it has been developed to cater for one particular product or similar shapes at a time. CaCODE should be easy enough to use so that it requires no prior CAD skills. For this reason, the first version of CaCODE was developed for pen design and was used to run trials in which members of the general public were asked to comment on whether or not it was easy to operate.Creating the pen designBefore creating the modifiable starting point 3D model pen, it was necessary to have an idea of what the pen should look like.

This stage required an experienced designer to produce a 2D sketch (Figure 4). Afterwards, a 3D CAD model was built in the Rhinoceros CAD system by interpreting the 2D sketches.242Figure 4: 2D sketches of the starting point pen designAfter generating the 3D model, it was then reconstructed by reducing the number of loft surface components, i.e. the profile curves, to a minimum (Figure 5). The profile curves could then be used by consumers as modifiable parameters. The number of curves determines the degree of freedom. The more parameters that can be changed by consumers, the more design freedom they have, but also the higher the level of difficulty. On the other hand, having fewer parameters will limit consumers' freedom in playing with the shape while presenting a lower level of difficulty.Figure 5: Simplifying the profile curves used for CAD constructionBuilding the CaCODE:Pen systemCaCODE:Pen was programmed using the Grasshopper software extension to Rhinoceros.

Grasshopper is a graphical algorithm editor integrated with Rhinoceros 3D modelling tools. This software creates a “programmable” CAD model but it requires no knowledge of programming or scripting. Although Grasshopper requires substantial effort in building the model definition, the result can be easily operated by non-designers to modify even very complicated 3D constructions [15]. With this feature, a design tool was created where consumers could modify the shape of the pen either by sliding buttons or by directly modifying the control points of the 3D model.Figure 6 shows how the CaCODE:Pen tool was generated from the Grasshopper software. Basically, the pen is built by four sections of circles where the top and middle circle positions are movable and the diameter of each circle is changeable. Also, by using another parameter, the circles can be modified to become ellipses. The bottom and end circles are fixed, because they are related to the ink refill.

Also, by using another parameter, the circles can be modified to become ellipses. The bottom and end circles are fixed, because they are related to the ink refill. The sliders presented in the top left box are the top circle parameters243and those in the bottom left box are for the middle circle parameters. The larger box on the right is the algorithm definition which processes the input from the sliders thus creates the surface of the pen.Figure 6: Grasshopper: Building a script for the CaCODE:Pen toolFigure 7 shows how the CaCODE:Pen interface appeared to the users. There is a control panel on the right and by using the mouse users can change the pen by sliding the small buttons on the sliders.

There is a control panel on the right and by using the mouse users can change the pen by sliding the small buttons on the sliders. In addition, the interface has 3 display views, front, right and isometric, to evaluate the appearance of the pen that they are designing.Figure 7: The user interface of the CaCODE:Pen: tool244Creating a range of AM samplesThis second part of this research was aimed at assessing how consumers would respond to the materials, surface finish, and colours of consumer products made using AM. To this end, 14 pen samples were built based on three varying parameters: different AM technologies and build materials, different colouring techniques and different finishing techniques (Figure 8). The AM technologies selected for this study were laser sintering/melting, polyjet printing, polymer extrusion and ink-jet printing onto powder.

The AM technologies selected for this study were laser sintering/melting, polyjet printing, polymer extrusion and ink-jet printing onto powder. Suitable post processing techniques were used to achieve better surface finish and colour quality, including tumbling, painting, dying and metallisation. Two of the pens were hand-finished and painted to provide injection moulded quality, both gloss and matt finish, so that they could act as datums against which the quality of the other pens would be evaluated.Figure 8: AM pen samples used for the studyUser TrialsThe next phase of the work was to present the software to a reasonably representative sample of the general population and ask them to design their own pens. Initially, ten participants were recruited, eight of whom had never used CAD before. Their ages ranged from late twenties to early forties and their backgrounds ranged from research fellows through to homemakers.

Their ages ranged from late twenties to early forties and their backgrounds ranged from research fellows through to homemakers. The participants were introduced to the software and asked to use it to create their own design of pen. They were then asked to evaluate their experience by completing a questionnaire. Research participants were also asked to evaluate the appearance of the range of AM made pens.

A summary of the questions asked in the questionnaire is as follows: • Have you ever wanted to design a product, if not, why not? • Have you ever actually designed a product, if not, why not?245If yes, give an example of a product that you have designed and state how you designed it.Have you ever wanted to create/modify a product, if not, why not? • Have you ever actually created/modified a product by yourself, if not, why not? • • Having now used PenCAD as consumer design software, how do you rate its usability? • Are there are any feature(s) of PenCAD that you feel are inconvenient? • Are there any feature(s) you would like added to PenCAD? • Having now seen the 3D printed products, considering the surface finish and the colours, could you please rank your 5 most preferred pens and your 5 least preferred pens.If yes, how did you create it?When considering which pen you would like to buy, how important is function, surface finish and colour quality?Having now used consumer design software, how much would you like to design a pen? • Having now seen pens built using 3D printing, how much would you like to buy one? • If you could design and/or print products for yourself, would you keep/use/maintain them longer than previous standard products that you have used?From your 5 favourite pens, please say what price would you pay for each of them?Many of the questions were asked in a way that enabled them to be answered using a 5 point Likert scale ranging from “strongly disagree” through to “strongly agree”.

The values given by the participants are shown as colour-coded bars in Figure 9, where each horizontal bar represents the ten answers. The most important results to note are that sliders were seen as the most useful interaction tool, that functionality and surface finish were more important than the colour quality, that downloading a ready-to-print design was slightly more desirable than designing one using PenCAD and that most of the participants thought they would keep a customised AM pen longer than a standard design. There is a tension between these last two results which will need to be investigated further.I found it is easy to change the shape using sliders I found it is easy to change the shape using click & dragUsing a ruler is helpful to change the shapeThe function of the product is importantThe surface finish of the product is importantThe colour quality of the product is importantI would like to design my own product using consumer design software (e.g.

PenCAD) I would like to download a ready-to-print product design I would like to buy my own product made from a 3D printer I would still prefer to buy a standard product (available on the market) I would keep/use/maintain them longer than previous standard products that I have usedSTRONGLY DISAGREEDISAGREE NEUTRAL AGREESTRONGLY AGREEFigure 9: Results of Likert scale questions.246The participants’ pen ranking values were inverted to give each pen a score out of 10 (i.e. a top ranked pen was given a score of 10 and a bottom ranked pen was given a score of 1) and used to calculate the total score given to each pen by the 10 participants. These scores are shown in Figure 10 next to a brief description of each pen, i.e. how it was made and finished. Not surprisingly, the two injection moulding quality pens came out top but several of the AM made and finished pens came quite close, specifically laser sintering with tumble and dyed finish and polyjet with tumbling.

It is interesting to note that the uncoloured untumbled laser sintered pen gained a higher score than the uncoloured tumbled version. In general, pens with a rougher finish or more visible layers seem to have come out worst. The only pen made in metal received the lowest score with several participants commenting that it was too heavy. In terms of the price that consumers said they would pay for personalised pens, almost all the pens were valued at between 1 and 5 GBP. Although such a price is more than that of a standard injection moulded pen, it would not justify the use of high-end AM machines.Figure 10: Scores out of 100 for the range of pens.Conclusions and Future WorkThe research has identified the potential for simplifying the interfaces of 3D CAD systems to a level where they can be used by non-expert users. The user of slider buttons was a particularly effective way of simplifying the modelling process.

The user of slider buttons was a particularly effective way of simplifying the modelling process. The general reaction to CaCODE:Pen was favourable although some consumers felt restricted by the shape of the starting pen design that was used. A possible solution to this would be to create a starting design that was built from sketches drawn by the consumer. This method has been used previously at Loughborough [16] but suffered from two drawbacks. Firstly, it was time consuming for the designer and secondly, some consumers felt uncomfortable with sketching. An alternative is to use software that will automatically generate a range of potential designs, from which the consumer can select a starting point. Such a system is “evoShape” (also developed at Loughborough University), which is an interactive evolutionary design (IED) tool that generates shapes through the use of the Genetic Algorithm [17].

Such a system is “evoShape” (also developed at Loughborough University), which is an interactive evolutionary design (IED) tool that generates shapes through the use of the Genetic Algorithm [17]. Figure 11 shows an example of evoShape247output where one design has been selected by the user for further development. Future work will investigate how IED tools can best support consumer design.Figure 11: Automatically generated designs in the evoShape systemThe research has also identified several AM systems and post-processing combinations that might be suitable for the manufacture of consumer products. In general, many of the AM parts evaluated were found lacking in some way by the consumers. However, laser sintering and polyjet technologies show good promise with participants in the trial scoring tumbled versions of these pens close to injection mould quality pens.

However, laser sintering and polyjet technologies show good promise with participants in the trial scoring tumbled versions of these pens close to injection mould quality pens. It was encouraging to note that most of the participants believed they would keep a personalised pen design longer than a standard design. This could be a route towards improved sustainability and away from the “throw away” society. However, the range of prices that consumers said they would pay for the pens was not high enough to justify the use of high-end AM systems, at least for this product. This could possibly be remedied in future by the use of entry-level 3D printers. Although the quality of parts from these machines is improving, previous research has shown that they are not yet capable of consistently producing high quality parts [18].

Although the quality of parts from these machines is improving, previous research has shown that they are not yet capable of consistently producing high quality parts [18]. Future research work will look at build strategies and finishing techniques that can be applied to entry-level 3D printers with a view to making them suitable for manufacturing of consumer designed products.AcknowledgementsThe authors gratefully thank Dr. Allan E.W. Rennie and all of the Lancaster Product Development Unit staff at Lancaster University for manufacturing some of the samples in SLM metal. Additional thanks are also extended to Laserlines for building some of the pens using FDM and to 3TRPD for tumbling the products.248References[1] M. Sinclair, “A Classification of Consumer Involvement in Industrial Design,” in DesignSchool Research Student Conference (DeSReS), Loughborough University, 2011.[2] H.

However, ensuring a consistent quality of parts processed by PBF-LB/M is crucial to compete with established manufacturing processes. In-situ process monitoring systems, such as coaxial melt pool monitoring (MPM), can contribute to this goal by minimizing post-process quality control. Three monitoring systems, a commercially available MPM system, an optical high-speed camera, and a thermal high-speed camera, were compared to identify process phenomena. Secondly, the suitability of the MPM system for in-situ quality control was tested by employing novel gas mixtures in the process. The mixtures include argon (Ar) with hydrogen (H2), helium (He), and carbon dioxide (CO2). The first results showed the capabilities of the MPM system to monitor relevant process anomalies.

The mixtures include argon (Ar) with hydrogen (H2), helium (He), and carbon dioxide (CO2). The first results showed the capabilities of the MPM system to monitor relevant process anomalies. Also, the addition of He and H2 to the process gas resulted in an improvement in the melt pool stability and a reduction of process by-products compared to Ar.IntroductionAdditive Manufacturing (AM) is used more frequently in the industry to produce complex parts [1]. This manufacturing technology is already the standard for special applications in the medical or aerospace industry [2, 3]. Compared to conventional manufacturing processes, AM offers the freedom of design and functional integration. This is essential to meet the increasing requirements of the industry in terms of parts and the cost-effectiveness of individual projects [4]. One of the most widespread AM technologies is a process called powder bed fusion of metals using a laser beam (PBF-LB/M) [5].

One of the most widespread AM technologies is a process called powder bed fusion of metals using a laser beam (PBF-LB/M) [5]. In this process, a metallic powder is distributed by a recoater (R) and melted locally layer-wise by the energy of a laser source [5]. A protective gas-shielded process chamber is crucial to ensure consistently high process stability and part quality. This atmosphere limits the oxidation and nitriding reactions of the powdered material and the melt pool. Dietrich et al. [6] stated that a minimum residual oxygen content of 1000 ppm (typically the lowest machine standard) is insufficient for processing titanium and its alloys. Pauzon et al. [7] supported this thesis and investigated various alloys and their part properties at residual oxygen contents below ten ppm. Another function of the process gas consists of the continuous recirculation of the gas above the build platform, ensuring the removal of process by-products from the interaction zone of the laser.

Another function of the process gas consists of the continuous recirculation of the gas above the build platform, ensuring the removal of process by-products from the interaction zone of the laser. Ladewig et al. [8] studied the influence of the gas flow speed and the resulting process stability and part quality. The authors indicated that a low gas speed leads to low laser coupling due to process by-products. Reijonen et al. [9] could confirm this phenomenon. Furthermore, the process gas has also been identified in various publications as a new individual process parameter in addition to1021the conventional laser parameters [10-14]. In PBF-LB/M, nitrogen (N2) or argon (Ar) are commonly used, depending on the material [15]. Novel process gas mixtures containing helium (He) or hydrogen (H2) in Ar are being studied. Bidare et al. [12] compared printing in Ar and He high-pressure atmospheres.

Novel process gas mixtures containing helium (He) or hydrogen (H2) in Ar are being studied. Bidare et al. [12] compared printing in Ar and He high-pressure atmospheres. Besides a reduced recoil pressure due to the high-pressure atmosphere, using He led to less formation of process by-products. Pauzon et al. [16] were able to increase productivity and part quality on Ti - 6Al - 4V by adding fractions of He to the process gas. Deckers et al. [11]identified a reduction in surface roughness and increased part density using alloy 718. Thus, a process influence can be generated via the type of protective gas, its velocity, its flow profile, or the chamber pressure. As a wide range of process influences exist, a constantly reproducible part quality must be ensured for an industrial application. This can be verified through process monitoring. A distinction is made between in-situ and subsequent monitoring. The present study focuses on in-situ monitoring.

This can be verified through process monitoring. A distinction is made between in-situ and subsequent monitoring. The present study focuses on in-situ monitoring. In-situ monitoring can be differentiated into on-axis and off- axis systems. On-axis systems like the commercially available EOSTATE melt pool monitoring (MPM) acquire emission data within the laser beam path via beam-splitting mirrors [17]. The acquired data depends on the implemented sensor type. For example, single-channel sensors such as photodiodes, pyrometers, or high-speed cameras are commonly used [18-20]. On the contrary off-axis monitoring captures process information at a fixed angle [19]. Thus, it can be used to validate on-axis data. MPM offers a high temporal (60 kHz) and spatial resolution (50 µm/pixel) [17]. However, little knowledge about the correlation between the digit sensor signal and the influence of process phenomena is available. Stutzman et al.

However, little knowledge about the correlation between the digit sensor signal and the influence of process phenomena is available. Stutzman et al. [20] investigated optical emissions during PBF-LB/M changing the gas flow speed. Felix et al. [21]correlated the MPM sensor signal with part defects but did not analyze the influences of the signal. Due to MPM spot size, the measurement is influenced by many different process phenomena, such as process by-products (spatter, fumes,..), the solidified material, the powder bed, and the melt pool. These influences are to be differentiated and analyzed with two off-axis setups within this paper. In addition, the influence of novel process gas mixtures will be assessed with MPM and two custom-built monitoring systems.Materials and MethodsPBF-LB/M ProcessThe experiments were carried out using an EOS (Electro Optical Systems GmbH, Germany) M290 PBF-LB/M system.

In addition to a melt pool monitoring system from EOS, custom-made monitoring systems were installed. The material used was alloy 718 with a 22 - 44 µm particle size distribution provided by Praxair Surface Technologies. The process parameters were optimized based on the commercially available alloy 718 EOS process parameters for a layer thickness of 40 µm. In order to meet the industry’s demand for increasing productivity, the study was conducted not only with the standard layer thickness of 40 µm but also with adapted process parameters and a layer thickness of 120 µm. Table 1 lists the utilized process gas mixtures with an excerpt of their physical properties. The gas speed of an EOS M290 is controlled by differential pressure in the central gas ductwork. Thus, the differential pressure had to be determined with respect to the physical properties of each process gas mixture. This ensures an equal and uniform gas flow and speed (G).

Thus, the differential pressure had to be determined with respect to the physical properties of each process gas mixture. This ensures an equal and uniform gas flow and speed (G). A build plate was equipped with several measuring points in the preceding study. These points were analyzed using different differential pressures utilizing a Testo 416 vane anemometer. The differential pressure for gases with a He-content higher than 50% had to be adjusted. Before each of the six print jobs, the chamber volume was opened and re-purged three times with the respective gas to ensure a full exchange of the process atmosphere. To safely use gases containing H2, the ADDvance® O2 Precision was used.

To safely use gases containing H2, the ADDvance® O2 Precision was used. This1022device analyzes and controls the oxygen and humidity level inside the process chamber and is unaffected by H2.Density in kg/m3 1.78 0.18 1.30 0.66 1.32 1.70Thermal Conductivity λ in W/(m*K) 0.02 0.15 0.04 0.08 0.04 0.02Specific heat capacity cp in J(kg*K) 0.52 5.19 0.71 1.41 0.76 0.57NameAr He Ar + 30%He Ar + 70%He Ar +30%He +10%CO2 Ar +5%H2 Table 1: Physical properties of the employed gases at norm conditions (p=1.013 bar, T=0°C).The job design consists of three cuboid specimens (10 x 10 x 10 mm). Three single scan tracks (SL) are on top of each specimen. The length of each SL is 10 mm.Process Monitoring Thermal High-Speed SetupInitially, the process of an EOS M290 can only be observed by the laser-safety window located at the machine's front door. This window contains a safety glass that filters the radiation of the laser beam at 1064 nm.

This window contains a safety glass that filters the radiation of the laser beam at 1064 nm. To utilize a thermal high-speed (THS) camera, a window with a different transmission wavelength band had to be used. Hence, an additional observation area was necessary. Therefore, an unused cable feedthrough at the top of the build chamber was modified. A specific infrared long-pass filter was chosen with a transmission wavelength band of λ= 2.5 - 4.8 µm to observe the process with a THS camera. The filter was placed in a custom-made aluminum mount, which fit into the cable feedthrough. Since the filter is not sufficient protection against reflection laser radiation, an additively manufactured 316L stainless steel cover was installed between the camera and the aluminum filter housing. The setup-up, highlighted in Figure 1, allowed safe monitoring of a circular area.

The setup-up, highlighted in Figure 1, allowed safe monitoring of a circular area. The approximate diameter is 50 mm near the gas inlet on the build platform.1023Infrared windowApronforlasersafetyOptical camera axisField of viewLight-tight sealingGR(a)(b)Thermal high-speedcameraxyzFigure 1: Schematic layout of the THS study (a) and final experimental setup (b).A FLIR X6900sc THS camera with a 100 mm focal distance lens was used. A 3/4 - inch spacer ring was used between the camera and the lens. This enabled to control the focus on the process zone. The camera was mounted on a tripod in front of the M290. To run the camera, the software FLIR ResearchIR 4 was utilized. This software facilitates camera control and allows for adjustments to various settings. The trials were recorded with and without a neutral density (ND) filter at a 320 x 256 pixels resolution to allow the camera’s maximum framerate of 1956 Hz.

The trials were recorded with and without a neutral density (ND) filter at a 320 x 256 pixels resolution to allow the camera’s maximum framerate of 1956 Hz. If FLIR superframing was used, the framerate was set to 862 Hz. Since the emissivity within the PBF/LB-M process is unknown and subject to ongoing research, an alternative approach was made to compare the monitored temperature. It was assumed that the emissivity coefficient within the study remains the same regardless of the process gas. Thus, the process gas is the only changing variable. A custom MATLAB-based script processed the videos, normalized them, and plotted them side-by-side. The temperature values are normalized using the upper and lower limits of the corresponding temperature range employed during a specific recording.High-Speed Setup Using a high-speed (HS) camera for process monitoring positioned in front of the laser safety glass at the front door is not ideal due to the resulting viewing angle.

Therefore, an existing recess on the right side of the build chamber was modified, in which an LED for the illumination is located. A fully transmissible glass plane was initially used to cover the original LED light source. This pane was exchanged for a laser safety glass equivalent to the one mounted in the process chamber door. A light source consisting of three 32 - Watt LED panels had to be installed in the build chamber near the ceiling. Especially for HS cameras, bright illumination is necessary due to the fast shutter speed. The additional light allowed an increase in the frame rate. The camera, a Photron FASTCAM NOVA S 12, and a deflection mirror were placed on a mounting rail to adjust the viewing angle and focus. A Nikon Micro-Nikkor 200 mm 4,0 D IF-ED lens was attached to the HS1024camera. The framerate was set to 4000 Hz with a shutter speed of 1/16000 seconds at a resolution of 1024 x 1024 pixels.

A Nikon Micro-Nikkor 200 mm 4,0 D IF-ED lens was attached to the HS1024camera. The framerate was set to 4000 Hz with a shutter speed of 1/16000 seconds at a resolution of 1024 x 1024 pixels. The recording was controlled with the Software Photron FASTCAM Viewer 4 and exported to MATLAB for further custom analysis. The HS setup, displayed in Figure 2, ensures an observation of the process perpendicular to the gas flow. After recording, the videos were further processed and plotted using MATLAB.xyHigh-speed cameraMirrorField of viewRecoater(a)(b)Single LinesRGFigure 2: Schematic layout of HS setup (a) and CAD mock-up (b).Evaluation for MPMA custom methodology was developed to gain a further understanding of the MPM signal. Firstly, influences on the sensor's signal data had to be investigated to make correlations between the signal data and the process phenomena.

Firstly, influences on the sensor's signal data had to be investigated to make correlations between the signal data and the process phenomena. Prior to the measurements, the EOSTATE MeltPool system had to be calibrated to ensure a precise measurement of the geometrical and intensity data. Preliminary studies were conducted to gain a deeper understanding of the process gas impact on the MPM. This preliminary study involved the recording of the MPM signal varying the laser power. Within the main study, only the process gas was changed. After each print job, the respective MeltPool raw data was exported. A workaround had to be used since the raw data could not be acquired from the EOS software. First, the data of each layer was exported as .fig files. This data includes the x & y-position of the laser focus spot, the on & off-axis intensity measured by the photodiode, the exposure type, the modulation signal, and the laser power.

This data includes the x & y-position of the laser focus spot, the on & off-axis intensity measured by the photodiode, the exposure type, the modulation signal, and the laser power. The data of each layer was then converted to a MATLAB-based matrix representation for further processing. A custom- made MATLAB APP was created to display, segment, analyze, and export the acquired data according to the requirement. The MPM, the MPM data, and the workflow are shown in Figure 3. Only a small part of the custom software was used for the present study.

With the ND filter activated, the recording was limited to high-temperature observations. Therefore, the melt pool could be analyzed as unsaturated. It was observed that increasing the He or H2 content led to a decrease in the melt pool surface temperature. This observation was made by comparing the maximum normalized temperatures in a 20-pixel region of interest (ROI) around the melt pool of the 40 µm layer thickness single lines. The results are displayed in Figure 4. This might be due to an increased heat capacity and conductivity of the novel gas mixtures compared to an Ar-atmosphere.10260204060801001200.610.70.80.9RelativetimeinmsArMaximum normalized temperatureAr+30%HeAr+70%HeHeFigure 4: Comparison of the normalized maximum temperature of the THS camera.Without the ND filter, the melt pool size, its cooling behavior, the process by-products, and their temperature profile could be made visible.

If one compares the synchronized frames, it is apparent that the process gas significantly influences the amount and temperature of process by-products. This effect is also present at a 40 µm layer thickness.010.5Temperature (normalized)ArAr+30%HeMeltingCoolingMeltingCoolingAr+70%HeHeAr+5%H2Ar+30%He +10%CO2Figure 5: Sequence of synchronized THS videos at the melting and cooling stage.1027This is already visible using a layer thickness of 40 µm. He- and H2-containing atmospheres create fewer process by-products and significantly reduce their temperature. Using pure He doesn’t even create high-temperature by-products, based on the pseudocolor image in comparison to pure Ar. This could be caused by a change in the recoil pressure or the thermal properties. Due to the fast cooling, particle adherence to solid surfaces is also greatly influenced. This correlates with the findings of Deckers et al. [11] for an improved surface quality using He- or H2-containing gases.

This correlates with the findings of Deckers et al. [11] for an improved surface quality using He- or H2-containing gases. As the layer thickness increased, the occurrence of process anomalies also rose. This effect is due to an increase in laser-powder interaction. There was a glowing behavior of the melt pool under the Ar-atmosphere notable. Within the 120 µm study, a balling effect (unstable, discontinuous melt pool tracks) for pure Ar Ar-H2 and low Ar-He mixtures could be identified. Mixtures with a He fraction above 30 Vol.-% created homogenous scan tracks even for 120 µm layer thickness. The influence of CO2-containing mixtures known from Gas Metal Arc Welding (GMAW) did not arise. No increased energy input or significant increase in process by-products was detected using CO2 containing mixture.

No increased energy input or significant increase in process by-products was detected using CO2 containing mixture. This could be due to the present helium content of that mixture.Meltpool Monitoring AnalysisGenerally, the MPM sensor signal data exhibits fluctuations within approximately 10 % of the mean value. The sensor signal is depicted in Figure 6 a). In order to gain a better understanding of the influence of process gases, various laser powers were examined using different process gases. This is evident in Figure 6 b), where higher laser powers correspond to an increase in MPM intensity. This effect is more pronounced at lower laser powers (PL) and increased the average MPM Intensity (ISMPM, avg). For instance, using an Ar atmosphere, increasing PL by 50 W from 135 W led to an average intensity increase of 882 digits, while the same increase from 285 W caused an average increase of 409 digits.

For instance, using an Ar atmosphere, increasing PL by 50 W from 135 W led to an average intensity increase of 882 digits, while the same increase from 285 W caused an average increase of 409 digits. At a PL of 285 W, the measured ISMPM avg is 4646 digits for an Ar atmosphere, 4637 digits for Ar +30 Vol-%, 4429 digits for Ar +70 Vol-%, and 4243 digits for He. The difference between the pure Ar and He atmosphere amounts to 403 digits, similar to the difference observed with a 50 W increase in PL. It should be noted that the values displayed in Figure 6 b) are obtained for single scan tracks conducted against the gas flow at 40 µm layer thickness.PLin W135185235285335370100030005000ArHeAr+30%HeAr+70%HeISMPM,avg. indigits7000(a)(b)10000510Lengthofscantrackinmm30005000Figure 6: Fluctuation of sensor data for a single scan track (a) and the comparison of the influences of process atmosphere and laser powder on ISMPM,avg.

(b).1028High-Speed AnalysisThe HS recordings highlighted, similar to the THS results, a change in the process by varying the process gases. First, an evaluation of the influence of incandescent by-products on the MPM signal was made. Therefore, the MPM intensity was plotted in addition to a frame-wise spatter count algorithm. From Figure 7, it is apparent that a correlation between the MPM signal and the occurrence of process by-products can be made. A significant reduction of signal strength and sensor fluctuation was present using He-containing gases. The Ar-atmosphere created more sensor fluctuation and by-products.0501000200040006000Ar+70%He050100020004000HeISMPM,avg. indigitsSpatter count in unitsRelativetimeinms02040608010012005010002000400060008000Ar0501000200040006000Ar+30%HeFigure 7: Comparison of ISMPM,avg.

indigitsSpatter count in unitsRelativetimeinms02040608010012005010002000400060008000Ar0501000200040006000Ar+30%HeFigure 7: Comparison of ISMPM,avg. and the spatter count using different process atmospheres.Additionally, if zoomed to a field of view (FOV) of the melt pool, the process fume could be investigated. The FOV of the employed process atmospheres are displayed in Figure 8. The He- enriched atmospheres seemed to have decreased fume density. Furthermore, the removal of fumes was affected by the process gas used. High He contents above 50 Vol.-% created swirls in comparison to a more straight laminar flow of Ar and Ar+30 Vol.-% He. This could be caused due to an insufficient flow speed. Interaction with the laser beam and the fumes were visualized for Ar and low He-containing gases. Since the fume density was visually analyzed, no direct correlation to the MPM signal was made. However, a decrease in MPM signal intensity and a less dense fume behavior are contradictory.

Since the fume density was visually analyzed, no direct correlation to the MPM signal was made. However, a decrease in MPM signal intensity and a less dense fume behavior are contradictory. Therefore, more intensity could reach the MPM sensor. Thus, the signal1029might not be influenced by the generated fumes. Low H2 and CO2 mixtures showed no significant influences. They were comparable to Ar +30 Vol.% He.GRGRGRGRArAr+70%HeAr+5%H2Ar+30%HeHeAr+30%He +10%CO2GRGRGRGRGRFigure 8: Image Sequence of HS videos to compare process phenomena using different process gases.ConclusionWithin PBF-LB/M, process stability plays a decisive role in a successful process and the associated part quality. This is challenged by requirements such as increasing productivity and the resulting change in layer thickness or, for example, the identification of new alloys.

This is challenged by requirements such as increasing productivity and the resulting change in layer thickness or, for example, the identification of new alloys. The present paper analyzed the process stability based on commercially available process monitoring (MPM) and custom monitoring such as THS and HS in an industrial-sized PBF-LB-M machine. He- or H2- containing gases offer possibilities to overcome process instabilities and increase part quality. The following conclusions about process monitoring within PBF-LB/M could be made:MPM can be used to analyze process stability. This could be validated via a THS setup. However, MPM does not offer the possibility to identify individual process influences due to the measurement spot size and the off-axis recording.1030The MPM signal correlates with laser parameters and the used process atmosphere. • A custom HS setup could be utilized to monitor the process perpendicular to the gas flow at the level of the process chamber.

• A custom HS setup could be utilized to monitor the process perpendicular to the gas flow at the level of the process chamber. In addition to the analysis of process by-products, this setup also enables the realization of a schlieren set-up.A custom THS setup was implemented to analyze melt pool behavior. This setup facilitates the analysis of melt pool temperatures after further research on the emission coefficient.The following conclusions about the influence of novel process gases within the PBF-LB/M process could be made. Using novel, He- or H2-mixtures results in:A significant reduction in the generation of spatter and fumes. • A decrease in MPM surface intensity during welding. This is comparable to a reduction of laser power by 50W.Less chance of adherence for hot process by-products. Consequently, the surface quality increases, and surface roughness decreases.Faster cooling of the solidifying melt pool surface.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA COMPARATIVE AND EXPERIMENTAL STUDY ON THE EFFECT OF HEAT TREATMENT CYCLES FOR PBF Ti6Al4VM. Karasoglu*, E. Yasaꝭ, E. Tan**, A. Yağmur‡Eskisehir Technical University, Eskisehir, Turkey ꝭEskisehir Osmangazi University, Eskisehir, Turkey ** ASELSAN Inc, Ankara, Turkey ‡ EOS GmbH, Krailling, Germany AbstractPowder bed fusion (PBF) presents the highest level of technological maturity and industrialization level for metallic materials among other Additive Manufacturing technologies. The advantages of high geometrical complexity, ability to produce internal cavities, reduced lead time and buy-to-fly ratio enables a wide range of application areas from aerospace to biomedical. Laser-PBF and Electron-PBF present different limitations and opportunities while they can both build from Ti6Al4V powder.

Laser-PBF and Electron-PBF present different limitations and opportunities while they can both build from Ti6Al4V powder. The performance of the E-PBF and L-PBF parts highly depends on the resulting microstructures and differs significantly due to various mechanisms such as preheating temperatures and processing environment. Moreover, the obtained material properties generally necessitate heat treatments for reducing residual stresses, enhancing mechanical properties and changing the microstructure.

Moreover, the obtained material properties generally necessitate heat treatments for reducing residual stresses, enhancing mechanical properties and changing the microstructure. This study aims to investigate the effect of the same heat treatment cycles on the E-PBF and L-PBF microstructure evolution and microhardness by a comparative experimental work with several combinations of exposure durations, temperatures and cooling rates.Keywords: Laser Powder Bed Fusion (L-PBF), Electron Powder Bed Fusion (E-PBF), Heat Treatment, MicrostructureIntroductionAdditive Manufacturing (AM), also known as 3D Printing, is a new group of technologies based on layerwise production enabling very complex part geometries in contrast to subtractive manufacturing processes. Powder bed fusion (PBF), is one of the 7 sub-groups of AM technologies according to ASTM classification[1].

Powder bed fusion (PBF), is one of the 7 sub-groups of AM technologies according to ASTM classification[1]. PBF methods, excluding Selective Heat Sintering, employ a laser or an electron beam to provide a high energy to melt or to sinter powder particles. All PBF processes include spreading of powder material over a base plate or previously solidified layers although different mechanisms may be employed in various processes. The most commonly used PBF processes for metallic parts are Laser-Powder Bed Fusion (L-PBF), also known as Selective Laser Melting or Direct Metal Laser Sintering/Melting and Electron-Powder Bed Fusion (E-PBF), also known as Electron Beam Melting. Although there are some similarities between L-PBF and E-PBF processes, they have some major differences leading to significant changes in the obtained part properties.

Although there are some similarities between L-PBF and E-PBF processes, they have some major differences leading to significant changes in the obtained part properties. In the E-PBF process, due to the presence of an electron beam, the process is carried out under vacuum and high preheating temperatures are applied in order to provide necking in between loose powder particles to avoid smoke formation during scanning. In the L-PBF process, whereas, the preheating temperatures are quite limited up to 200 °C in most applications1076and the process is carried out under a protective atmosphere supplied by argon or nitrogen flow. The main differences between two processes are summarized in Table 1.PBF of materials lead to many advantages compared to conventional manufacturing. First of all, the waste material related to subtractive machining in the form of machining chips is very much reduced since the manufacturing route is additive.

First of all, the waste material related to subtractive machining in the form of machining chips is very much reduced since the manufacturing route is additive. This significantly reduces the buy-to-fly ratio in aerospace applications dealing with expensive materials. Moreover, very complex geometries including internal channels, cavities or lattice structures are possible with a relatively good dimensional accuracy. Moreover, using PBF, it is possible to produce mono-parts rather than assemblies reducing the need for joining processes and lead time.

This may also lead to increase in the life-time of the parts since joining processes such as welding are generally limiting factors in terms of life assessment.Table 1: Comparison of E-PBF and L-PBF processesProperty Energy Source Beam Size [2] Maximum Build Volume [3] Layer Thickness Minimum Wall Thickness [4] Build rate [cm3/h] [5] Dimensional Accuracy [4] Surface Roughness (Ra) [5] Geometry LimitationsAvailable MaterialsE-PBF Electron 0.2-1 mm 350 x 380 mm (Ø/H) (Arcam Q20plus) 50-100 µm 0.6 mm80 ±0.3 mm 20-30 µm Powder sintering due to high preheating temperatures in internal features Only conductive materials: Ti alloys, Titanium Aluminides, CoCr alloys, Inconel 718L-PBF Laser 0.1-0.5 mm 800 x 400 x 500 mm (Concept Laser X Line 2000R) 20-100 µm 0.2 mm5-20 ±0.1 mm 5-15 µm Support structures are needed for overhang surfaces with low angles Stainless steels, tool steels, maraging steel, nickel superalloys, Ti and its alloys, CoCr alloys, precious metals, Al alloys, Cu alloysAs shown in Table 1, titanium alloys can be processes by both PBF methods.

Ti6Al4V alloy is the mostly addressed Ti-alloy in AM due to its superior biocompatibility, excellent corrosion resistance, high specific strength, high melting temperature and high fracture toughness [6]. It is an α + β titanium alloy saving weight for highly loaded structural parts, especially in aerospace applications. Although this alloy is widely utilized in a wide range applications from marine, biomedical to chemical industries, some properties of Ti6Al4V alloys make it difficult to process including low thermal conductivity, high reactivity and its sensitivity to strain hardening [7-10]. Thus, it is difficult and costly to machine this alloy making AM a good choice [11-13].In PBF processes, generally different microstructures are obtained due to high cooling rates leading to different mechanical properties in comparison to cast and wrought materials and Ti6Al4V is no exception.

Figure 1 shows the phase diagram and phase transformation as a function1077of cooling rate for Ti6Al4V. As shown, the formation of α′ martensite necessitates a specific thermal history with high cooling rates along with build temperatures lower than the martensite start temperature (Ms). This martensite start temperature depends on the initial microstructure, composition homogeneity and impurity levels ranging from 575°C to 800 °C. A complete α′ martensite can form for cooling rates greater than 410 °C/s whereas low cooling rates smaller than 20 °C/s will not form α′ martensite.Figure 1: (a) Phase diagram of Ti6Al4V (b) Phase transformation as a function of cooling rates [6]In the literature, L-PBF and E-PBF processes have been shown to form different microstructures.

L-PBF is related to fine needle-shaped α′ martensite, with a typical lath width about 0.2–1 μm whereas for E-PBF leads to α + β dual phase, where the α-lath has an average length and thickness of 1–3 μm and 0.5–1 μm, respectively while β platelets are about 0.1 μm thick [6]. This difference in the microstructures between L-PBF and E-PBF processes is attributed to the thermal histories due to different preheating temperatures. The high preheating temperature in E-PBF acts as an in-situ heat treatment of α′ martensite and decomposition into α + β microstructure [14-17]. With a similar approach, it is also possible to produce α + β microstructure in as-built L-PBF parts as well by using a low laser power and a low scan speed to enable a long material-laser interaction time and high energy input [18].After PBF of Ti6Al4V, heat treatments are applied for reducing residual stresses, enhancing mechanical properties and varying the microstructure of the produced parts.

Various researchers have focused on the effect of heat treatments on different aspects such as mechanical properties [19, 20], fatigue performance [21-24], dimensional accuracy, corrosion resistance[25], biocompatibility [26], creep behavior [27], etc. In comparison to L-PBF, the number of studies on the heat treatment of E-PBF is much lower[11]. Moreover, there are very few number of studies investigating the effect of different heat cycles on the obtained microstructures from L-PBF and E-PBF processes. Therefore, this study aims to reveal and understand the effect of soak duration,1078temperature and cooling rate in a heat treatment cycle on the microstructure of as-built E-BPF and L-PBF in a comparative way.Experimental MethodologyIn order to understand the effect of different heat treatment cycles on the PBF parts, the specimens were first produced.

The L-PBF specimens were produced on an EOS M290 machine under argon environment with a layer thickness of 40 µm whereas the E-PBF parts were produced on an ARCAM A2X machine in vacuum with a layer thickness of 50 µm. The process parameters for both machines are the standard ones which were optimized for maximum density.After the productions were complete, the parts were cut off from the base plates and heat treated in Protherm PLF 130/6 furnace under ambient atmosphere in order to understand the oxide layer formation on the surfaces and microstructure evolution in the core of the part.

The test plan with different heat cycles is presented in Table 2.Table 2: Test Plan1 2 3 4 5 6 7 8 9 10 11 12 13Temperature (°C) Duration (h) 1020 1020 940 940 900 850 850 850 800 800 800 800 7301 2 2 2 2 2 5 8 2 2 4 4 2Cooling Rate Furnace Cool Furnace Cool Air Cool Furnace Cool Furnace Cool Furnace Cool Furnace Cool Furnace Cool Air Cool Furnace Cool Air Cool Furnace Cool Furnace CoolAfter the heat treatments were applied, the specimens were cut showing the planes parallel to the build direction and mounted in a resin and polished to a mirror finish. In order to reveal the microstructures, a solution of 91 ml water, 6 ml HNO3 and 3 ml HF (Kroll’s reagent) was used as the etchant and applied to surfaces. For quantitative analysis of microstructural images, image analysis was carried out by using FIJI software package. Different phases (α and β) in the OM images were classified by Trainable Weka Segmentation Fiji plugin.

Different phases (α and β) in the OM images were classified by Trainable Weka Segmentation Fiji plugin. For classifications, a random forest classifier with 2 random features, 200 decision trees, 2 decimal places for computational accuracy, and 5 image filters (Gaussian blur, Hessian matrix, Sobel, Gaussian distribution difference, and membrane projection) was utilized. Subsequent to image classifications, segmentation processes were applied to α phases in the images in order to evaluate individual α particles. Segmentations were performed with Disconnect particles” Xlib plugin for FIJI using a modified watershed algorithm. In this method, particles are disconnected according to predefined1079constraint proportionality constant (k=0.7) [28].

In this method, particles are disconnected according to predefined1079constraint proportionality constant (k=0.7) [28]. Image analysis was limited to the heat treated samples at 850°C or higher temperatures due to the resolution limitation of OM.Results and DiscussionAs-built microstructuresFirst of all, as-built microstructures coming from L-PBF and E-PBF processes have been investigated. The microstructures are presented in Figure 2. As evident, while L-PBF sample exhibits a martensitic microstructure α′ acicular grains, E-PBF sample reveals lamellar α+β basketwave (Widmanstätten) structure. Both microstructures are identical to manufacturing method.

Both microstructures are identical to manufacturing method. It is well known that difference between those two microstructures is due to higher solidification rate and lower process chamber temperature of L-PBF in comparison to E-PBF [29].a) L-PBFb) E-PBF128 µm128 µmFigure 2: As-built microstructures from a) L-PBF b) E-PBFEffect of annealing temperatureThe annealing temperature was varied between 730 and 1200 °C in this study in order to understand the evolution of the microstructure from L-PBF and E-PBF processes in a wide range of temperatures as demonstrated in Figure 3. While the temperature was changed, the time and cooling rate were fixed.

While the temperature was changed, the time and cooling rate were fixed. All images in Figure 3 are belong to furnace cooled samples and annealing temperature was fixed to 2h.1080@730 °C 2h@800 °C 2h@850 °C 2h@900 °C 2hL-PBF128 µm128 µm128 µm128 µm1081E-PBF128 µm128 µm128 µm128 µm@940 °C 2h128 µm128 µm@1020 °C 2h128 µm128 µmFigure 3: Effect of the annealing temperatureThe results show that annealing temperature is a very important parameter to alter the microstructures for both L-PBF and E-PBF samples. It can be seen from the Figure 3 that lower temperatures (<900°C) have minor effect on microstructures and almost no visible difference from the as-built samples was observed with OM. Moderate annealing temperatures (900°C and 940°C) which are below the beta transus temperature modified microstructures to some extent.

Moderate annealing temperatures (900°C and 940°C) which are below the beta transus temperature modified microstructures to some extent. It can be seen that α lath thicknesses of E-PBF samples were increased by heat treatments with moderate temperatures and fully α’ structure of the L-PBF samples were converted to α+β two-phase microstructure. However, columnar parent β structures were not modified. On the other hand, high temperature (above transus temperature) annealing induced major changes in the microstructures of both E-PBF and L-PBF samples. High temperature annealing transformed parent β columns to uniform equiaxed grains and needle-like α laths to globular α grains.Effect of annealing durationIn order to test of the annealing duration, heat treatments were applied at 850 °C for 2, 5 and 8 hours with furnace cooling. Moreover, tests were carried out at 1020 °C for 1 and 2 hours.

Moreover, tests were carried out at 1020 °C for 1 and 2 hours. The results are depicted in Figure 4 and Figure 5.1082L-PBFE-PBF@850 °C 2h128 µm@850 °C 5h128 µm@850 °C 8h128 µmGrain SizesFigure 4: Effect of the duration at maximum temperature of 850 °C for furnace cooling1083128 µm128 µm128 µmL-PBFE-PBF@1020 °C 1h128 µm128 µm@1020 °C 2h128 µm128 µmGrain SizesFigure 5: Effect of the duration at maximum temperature of 1020 °C for furnace coolingThe obtained results lead to the fact that increasing annealing duration increases size of α grains. For the samples annealed at 850°C, increasing annealing time from 2h to 5h slightly increased grain sizes and grain morphology was maintained. However, the samples annealed for 8h displayed locally globular grains in the microstructure. In high temperature heat treatments, the effect of the heat treatment time is more evident on the grain size and shape of the α grains.

In high temperature heat treatments, the effect of the heat treatment time is more evident on the grain size and shape of the α grains. Figure 5 shows that increasing annealing time from 1 to 2h, grain size of the samples were nearly doubled and almost all α grains were transformed to a globular shape.Effect of cooling rateAs shown in Table 2, the effect of cooling rate was tested at 940°C and at 800°C for 2 and 4 hours. The obtained microstructures for L-PBF and E-PBF processes are presented in Figure 6 and Figure 7, respectively.1084Air CoolingFurnace Cooling@940°C 2h128 µm128 µm@800°C 2h128 µm128 µm@800°C 4h128 µm128 µmFigure 6: Effect of the cooling rate for L-PBF partsThe results reveal that cooling rate has no remarkable influence on microstructure except for the samples annealed at 940°C.

Furnace cooling led to increased α lath thickness and formation of globular α colonies locally in the microstructure of the samples annealed at 940°C due to slow cooling rate than air cooling.1085Air CoolingFurnace Cooling@940°C 2h128 µm128 µm@800°C 2h128 µm128 µm@800°C 4h128 µm128 µmFigure 7: Effect of the cooling rate for E-PBF partsOxidation and Hardness ResultsTitanium and its alloys are easily oxidized when exposed to elevated temperature and oxyen bearing conditions. Since the heat treatments were carried out in the open-air atmosphere in this study, the oxide layer formations were examined by OM images obtained from the edge of the heat treated samples. The overall oxide reaction consists of oxide formation and inward oxygen diffusion into the bulk material. This diffusion leads to an oxygen enriched layer, which is continuous, hard and brittle named as alpha-case layer, beneath the oxide scale.

This diffusion leads to an oxygen enriched layer, which is continuous, hard and brittle named as alpha-case layer, beneath the oxide scale. In this study, as expected, all heat treated samples are covered by an oxide layer. It is shown in the Figure 8 and 9 that increasing temperature and increasing duration led to a thicker oxide layer on the surface of the samples.1086@800°C 2h FC@800°C 4h FC@850°C 2h FC@850°C 8h FC@1020°C 2h FCL-PBF128 µm128 µm128 µm128 µm128 µmFigure 8: Oxide layer images of the heat treated samples1087E-PBF128 µm128 µm128 µm128 µm128 µmFigure 9: Oxide layer thicknesses of heat treated samples a) L-PBF b) E-PBFIn Figure 10, hardness results show that supertransus temperature annealed samples displayed extremely high hardness values (up tp 500 HV) despite of high grain sizes in contrast to average hardness around 380 HV.

Samples with larger grain sizes are normally expected to exhibit lower hardness values due to the Hall-Petch grain size effect [30]. Calculated grain sizes of the heat treated samples by image analysis are exhibited in Figure 11 and it can be seen from the Figure 11 that supertransus temperature heat treated samples have the highest grain sizes. It is important to note that the scatter of the grain size measurements increases as the grain size is increased especially at elevated temperatures. This is attributed to the measurement method and needs further investigations. Extremely high hardness vales at elevated temperatures are thus not due to the grain size effect and attributed to oxygen diffusion inside the samples and hardening of the samples.

Extremely high hardness vales at elevated temperatures are thus not due to the grain size effect and attributed to oxygen diffusion inside the samples and hardening of the samples. In order to reveal the exact effects of heat treatment conditions on hardness, heat treatments will be conducted under a protective atmosphere as the future work.1088a)Figure 10: Hardness of the heat treated samples a) L-PBF b) E-PBFConclusionsIn this study, various heat treatment cycles were applied to L-PBF and E-PBF samples inorder to reveal the effect the annealing temperature, duration and cooling rate in a wide range. The main highlights of the study can be summarized as:From as-built microstructures, it is observed that while L-PBF sample exhibits a martensitic microstructure α′ acicular grains, E-PBF sample reveals lamellar α+β basketwave (Widmanstätten) structure.The annealing temperature was varied between 730 and 1200 °C in this study.

While lower temperatures (<900°C) have minor effect on microstructures and almost no visible difference from the as-built samples was observed while moderate annealing temperatures (900°C and 940°C) below the beta transus temperature modified microstructures to some extent. High temperature annealing transformed parent β columns to uniform equiaxed grains and needle-like α laths to globular α grains.1089Increasing annealing duration increases size of α grains. The effect of duration is more pronounced at elevated temperatures.The results also reveal that cooling rate has only remarkable influence on microstructure for elevated temperatures.Heat treatments under ambient atmosphere led to a thin oxide scale layer and alpha-case layer. The diffusion of oxygen inwards the part greatly increased the microhardness values.Figure 11: Grain size of the heat treated samples a) L-PBF b) E-PBF1090References[1] I.

The diffusion of oxygen inwards the part greatly increased the microhardness values.Figure 11: Grain size of the heat treated samples a) L-PBF b) E-PBF1090References[1] I. Astm, ASTM F2792-10: standard terminology for additive manufacturing technologies, ASTM International (2010). [2] A. Agapovichev, A. Sotov, V. Kokareva, V. Smelov, Possibilities and limitations of titanium alloy additive manufacturing, MATEC Web of Conferences, EDP Sciences, 2018, p. 01064. [3] Additive Manufacturing Machines, manufacturing/machines. (Accessed June 18,2021. [4] M.G. Martinez, Metal Additive Manufacturing & Case Studies, https://www.slideshare.net/Prizztech/metal-additive-manufacturing-manufacturing. June 6, 2021. [5] C.E. Olivier Rigo, Selective Laser Melting versus Electron Beam Melting, 2013. https://www.slideshare.net/carstenengel/selective-laser-melting-versus-electron-beam-melting. (Accessed June 18, 2021. [6] S. Liu, Y.C.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCOMPARATIVE COSTS OF ADDITIVE MANUFACTURING VS. MACHINING: THE CASE STUDY OF THE PRODUCTION OF FORMING DIES FOR TUBEBENDINGB. Previtali *, A.G. Demir*, M. Bucconi*, A. Crosato†, and M. Penasa†Department of Mechanical Engineering, Politecnico di Milano, Milan, 20156 Italy †BLM Group, Cantu, 22063 ItalyAbstractAdditive manufacturing processes for metallic components become economically viable when they substitute conventional processes that make use of moulds and dies to produce casting or semi-finished parts with high added value. Common examples are: i) components in aerospace or energy sectors obtained by investment casting in high-temperature alloys; ii) personalized prostheses and implants in biocompatible metals in the biomedical sector.

In both cases, the annual batch size is low and often limited to a single or few pieces. However, for many other sectors decision making for process substitution from conventional to AM processes requires a correct economic analysis. The cost of AM processes depends also on the use of technological advantages. The paper explores the economic feasibility of selective laser melting (SLM) process when producing tube-forming tools. The analysed industrial case addresses the whole annual production of bending tools, traditionally made by milling from a solid block.

The analysed industrial case addresses the whole annual production of bending tools, traditionally made by milling from a solid block. The aim of this work is to identify the levers of the process that make additive production advantageous, even in more traditional sectors like tooling when different tool materials are used (namely a tool steel and a bronze alloy) and when hybrid manufacturing (subtractive plus additive) is carried out.1IntroductionMetal additive manufacturing technologies provide several advantages in terms of shape and details obtainable on the single product. This enables highly customizable production and development of new products taking advantage of such geometrical flexibilities. Biomedical implants are an example to where highly customized production is relatively easy to justify in terms of cost. Weight reduction and in-process assembly of diverse components is another factor to the adoption of these techniques in the aerospace and aviation sectors.

Weight reduction and in-process assembly of diverse components is another factor to the adoption of these techniques in the aerospace and aviation sectors. However, metal additive manufacturing processes in general, and selective laser melting (SLM) in particular, are highly costly production options compared to more conventional ones. The high capital cost of the machines and low productivity rates are the main reasons for the high production costs. Indeed, when metal additive manufacturing processes are required to be evaluated as substitutes to existing manufacturing routes composed of conventional processes, an economical evaluation is of vital importance. Tooling applications possess peculiarities, such as low number components produced with a large variety of shapes and sizes, that might suggest the substitution of the conventional process route with an additive manufacturing (AM) process chain.

However, although the premises may appear highly suitable for AM, economic conditions may not necessarily be competitive to justify the substitution of conventional machining processes.Several works have been reported in literature dealing with cost modelling of additive manufacturing (AM) processes [1]. Indeed, these models have been developed to confront AM2816technologies with conventional ones. Commonly, they conclude that small lot production of highly complex parts is more convenient with AM processes compared to casting or plastic deformation, which are post-processed to the final shape. The absence of dedicated tools and moulds within the AM process is a determining factor for reduced costs. On the other hand, such a comparison between the AM processes and machining for the production of small lot productions has not received the same amount of attention.

On the other hand, such a comparison between the AM processes and machining for the production of small lot productions has not received the same amount of attention. Within a conventional production route, machining processes can give net-shape form to simple feedstocks such as bars, plates, and blocks by the use of standard tools. The production of tube bending tools is such a case whereby a small lot of highly variable shapes are produced by conventional machining processes.Hopkinson and Dickens [2] proposed a cost model for AM processes considering three distinct technologies namely stereolithography, fused deposition modelling (FDM), and laser sintering (LS). The proposed model was simplified as it considered a limited number of input parameters. A main disadvantage is related to the fact that the model considered a single component throughout the whole production, leaving out a key advantage of the AM processes related to diversified production.

In particular, the model did not consider the general costs, while it has been demonstrated that they tend to have a higher impact on the final cost as the level of production automation increases [3]. Despite considering a post-processing stage, the model did not include any preparatory phase, which might have a marked impact especially when customized products are concerned. The post-processing stage is considered only in terms of manual labour, whereas the use of finishing machinery is not considered.Ruffo et al. [3] introduced a model for laser sintering, which can be easily adapted to SLM, due to the similarities between the two technologies. Also in this case, the model is rather simple to implement, yet lacks consideration regarding the post-processing stage, general costs and considered a single component throughout the production. Later on, Ruffo and Hague [4] introduced a modified model considering mixed production. More recently, Rickenbacher et al.

Later on, Ruffo and Hague [4] introduced a modified model considering mixed production. More recently, Rickenbacher et al. [5] proposed a cost model specific to the SLM process. Differently, this model considered a decomposition of the process in different phases, considering the associated cost item. The model also considered the details regarding pre- and post-processing as well as consumables such as process gas and material change. General costs and considerations over mixed production are missing, leaving a margin for improvement for future models. Lindemann et al. [6] also proposed similar improvements. However, also in this case mixed production has been left out and the cost of a single component type has been analysed.In this work, a new cost model specific for additive manufacturing by selective laser melting is proposed. The case study consisted of the production of customizable tube bending tools with high variability in shape and limited lot over the year.

The case study consisted of the production of customizable tube bending tools with high variability in shape and limited lot over the year. The novel aspects of this model are as follows: • Cost model considering all the production phases (i.e. pre-, AM, post-production). • The use of a real industrial case comparing the results to the existing conventional production performance. Comparison with a challenging conventional production case with machining, where dedicated tools are not required and small lots are considered.Model applied at build job level with mixed products.

•Considering part modifications such as hybrid production and weight reduction enabled by the SLM process.2817Table 1 – Nomenclature, definition and units of the terms in the cost modelsSymbol CT,P C1 C2 C3 C4 CT,Add CT,Pre CT,SLM CT,Post Cb,Add Cb,Pre Cb,SLM Cb,PostDefinition Total cost for a generic production process per year Material cost Tooling and consumables cost Equipment cost Overhead cost Total cost for the additive SLM production process per year Total pre-processing cost Total SLM production cost Total post-processing and finishing cost Total additive SLM production cost per build-job Total pre-processing cost per build-job Total SLM production cost per build-job Total post-processing and finishing cost per build-jobSLM PRODUCTION COSTCb,SLM,1 Cb,SLM,2 Cb,SLM,3 Cb,SLM,4 Cm Cre Cplate Cgas cyl Cfilter CC Ċoh Lre Lplate Lgas cyl Lfilter Vp Vb VW f rsol rpow N L two tbMaterial cost in SLM process per build-job Tooling cost in SLM process per build-job Equipment cost in SLM process per build-job Overhead cost in SLM process per build-job Material cost Cost of the re-coater blade Cost of the substrate plate Cost of the argon canister Cost of the filter cartridge Capital cost of the SLM system Overhead rate for the SLM operations Re-coater blade life per build job Substrate plate life per build job Argon canister life per build job Filter cartridge life per build job Volume of the part in the build-job Volume of the build-job (print area x height of highest part) Volume of un-melted powder in the build-job Not-recycled powder fraction Solidified material density Powder material density Number of parts produced per build-job Load factor of SLM system Capital write-off time for SLM system Building time per build-jobPRE-PROCESSING PRODUCTION COSTCb,Pre,3 Cb,Pre,4 Ccomp tcomp two,comp Lcomp Ċoh,compEquipment cost in the pre-processing per build-job Overhead cost in the pre-processing per build-job Capital cost of the hardware and the software licences Computing time per build-job Capital write-off time for hardware and software licences Load factor of the computing equipment Overhead rate for the computing operationCb,Post,3 Cb,Post,4 Cmill CEDM Coven Lmill LEDM Ċoh,fin tmill tEDMPOST-PROCESSING PRODUCTION COST Equipment cost in the post-processing per build-job Overhead cost in the post-processing per build-job Cost of the 5-axis milling machine Cost of the wire-EDM machine Cost of the induction heating furnace Load factor of the 5-axis milling machine Load factor of the wire-EDM machine Overhead rate for finishing operations Milling time per build-job Wire-EDM time per build-job2818Units € € € € € € € € € € € € €€ € € € €/kg € € € € € €/h # # # # mm3 mm3 mm3 % kg/ mm3 kg/ mm3 # % h h€ € € h h % €€ € € € €/kg % % €/h h h2 Cost model The cost model applied in this work is based on the classification of cost terms proposed by Ashby et al.

[7], where the total cost per each phase of a generic production cycle CT,P is composed by four terms (as shown in Table 1):𝐶",$ = 𝐶& + 𝐶( + 𝐶) + 𝐶*Equation 1with C1 material cost, C2 tooling and die cost, C3 equipment cost and C4 overhead cost. Usually, C1 accounts for all the materials used and wasted in the process, C2 accounts for all the tools and dies dedicated to the production of the entire batch, C3 accounts for the non-dedicated equipment and machine cost and C4 accounts for the overhead costs including direct costs (i.e. operator and energy), and indirect cost (i.e. administration, maintenance and safety).In the additive manufacturing with the SLM process the total cost of a product is the result of three fundamental steps: pre-processing, selective laser melting and post-processing. Therefore, the total cost CT,Add is given by three contributions:𝐶",+,, = 𝐶",$-.

Therefore, the total cost CT,Add is given by three contributions:𝐶",+,, = 𝐶",$-. + 𝐶",/01 + 𝐶",$234Equation 2where CT,Pre, CT,SLM and CT,Post are pre-processing, SLM production and post-processing costsrespectively.Since the SLM process is usually employed for the production of parts with different geometries and shapes in few quantities within a single build-job, it is convenient to evaluate the costs considering the build-job consisting of different parts rather than a single part. Accordingly, the total costs per build-job Cb,Add of all the parts that printed on the same platform is: 𝐶5,+,, = 𝐶5,$-.

Accordingly, the total costs per build-job Cb,Add of all the parts that printed on the same platform is: 𝐶5,+,, = 𝐶5,$-. + 𝐶5,/01 + 𝐶5,$234Equation 3where Cb,Pre, Cb,SLM and Cb,Post are pre-processing, SLM production and post-processing total costsper build-job respectively.2.1	 SLM	production	cost	per	build-job.	 The SLM production cost per build-job Cb,SLM, which defines the cost contributions related tothe AM production step can be formulated as:𝐶5,/01 = 𝐶5,/01,& + 𝐶5,/01,( + 𝐶5,/01,) + 𝐶	5,/01,*Equation 4where the single terms are:Material cost per build-job Cb,SLM,1In the SLM process, the material cost includes the powder used to selectively melt the volume of the printed N number of parts in the build job along with their supports and the un-melted powder (which is partially wasted and partially re-used).

Cb,SLM,1 can be further defined as:𝐶5,/01,& = 𝑉$ ∙ ρ32: ∙ 𝐶; + 	 𝑓 ∙ 𝑉= ∙ ρ>2? ∙ 𝐶;Equation 5where VP is the volume of all the parts and supports in the same build-job, rsol and rpow the density of solidified material and powder respectively, Cm the powder cost, VW the volume of the un-melted powder, f the powder fraction which cannot be recycled. The volume of the un-melted powder VW depends on the build-job volume Vb:𝑉? = 𝑉5 − 𝑉$Equation 6where Vb is given by the overall print area multiplied by the height of the tallest part of the build- job.2819Tooling and consumables cost per build-job Cb,SLM,2Usually Ashby’s model includes in the category C2 the cost for tooling and dies (which is affected by batch size and cost of tools/dies). However, in SLM production no moulds or dies are used while auxiliary tools and materials are consumed during each build-job, sometimes partially, sometimes totally.

However, in SLM production no moulds or dies are used while auxiliary tools and materials are consumed during each build-job, sometimes partially, sometimes totally. For every build job the substrate plate must be partially milled and the re-coater blade deteriorates. Moreover, per each job, a fixed quantity of inert gas (argon) is used and the filter cartridge is consumed. Therefore, Cb,SLM,2 is re-interpreted as the consumables cost per build-job:𝐶5,/01,( =𝐶-. 𝐿-.+𝐶>:B4. 𝐿>:B4.+𝐶CB3	DE: 𝐿CB3	DE:+𝐶FG:4.- 𝐿FG:4.-Equation 7where Cre, Cplate, Cgas cyl and Cfilter are the cost of the re-coater blade, plate, argon cylinder and filter cartridge while Lre, Lplate, Lgas cyl and Lfilter the respectively life expressed in numbers of build-jobs.

Equipment cost per build-job Cb,SLM,3The equipment cost per build-job Cb,SLM,3 can be estimated from the capital cost of theequipment CC :𝐶5,/01,) =𝐶H 𝐿 ∙ 𝑡?2𝑡5Equation 8once the building time per job tb, the load factor L (that means the time the machine isoperating) and the capital write off time two are known.Overhead cost per build-job Cb,SLM,4This includes the direct cost of the machine operator in addition to the indirect cost associated with floor space, energy and maintenance. In this model administrative costs (safety, administration, legal) are not considered. Hence, the overhead cost Cb,SLM,4 is estimated as:𝐶5,/01,* = 𝐶2J ∙ 𝑡5Equation 9where Ċoh is the overhead rate.2.2	 Pre-processing	cost	per	build-job	 The pre-processing phase is mainly constituted by operations at the computer for elaborating the 3D model, adding supports and applying design for additive rules along with post-processing of tool path and machine instructions.

In the present model, the pre-processing cost Cb,Pre accounts only for two main terms as:𝐶5,$-. = 𝐶5,$-.,) + 𝐶	5,$-.,* The single terms can be further defined as follows.Equation 10Equipment cost per build-job Cb,Pre,3Cb,Pre,3 is equipment cost of computer and software licences for CAD and CAM processing:𝐶5,$-.,) =𝐶D2;> 𝐿D2;> ∙ 𝑡?2,D2;>𝑡D2;>Equation 112820where Ccomp is the hardware and software cost, Lcomp and tcomp are respectively the loadfactor and time for computing the job file.

Overhead cost per build-job Cb,Pre,4The overhead costs per part is due to the operator cost, floor space, energy andmaintenance cost and can be estimated as:𝐶5,$-.,* = 𝐶2J,D2; ∙ 𝑡D2;>Equation 12where Ċoh,comp is the overhead rate.2.3	 Post-processing	cost	per	build-job	 The post-processing cost per build-job Cb,Post mainly accounts for three operations: 1) part removal from the substrate plate by means of wire EDM (along with support removal); 2) finishing of functional features and surfaces by milling; 3) stress relief and heat treatments.Cost of consumable material (lubricant and tool inserts mainly) could be neglected while no moulds and dies are used. Therefore, the tooling cost C2 is neglected and the post-processing cost is reduced to:𝐶5,$234 = 𝐶5,$234,) + 𝐶	5,$234,*Equation 13The single terms can be further defined as follows.Equipment cost per build job Cb,Post,3Cb,Post,3 is the cost of the equipment used for the removal and finishing operations (i.e.

wireEDM, milling and induction heating oven):𝐶5,$234,) =𝐶;G:: 𝐿;G:: ∙ 𝑡?2𝑡;G:: +𝐶KL1 𝐿KL1 ∙ 𝑡?2𝑡KL1 + 𝐶2M.N ∙ 𝑉$ ∙ 𝜌32:Equation 14where Cmill and CEDM are the cost of the wire EDM system and 5-axis milling center; with their respective load factors (Lmill and LEDM) and operation time per build-job (tmill and tEDM). Furthermore, Coven is the cost of the induction-heating oven per unit of mass treated. Overhead cost per build-job Cb,Post,4The overhead cost Cb,Post,4 is linked to the operator , floor space and energy cost and iscalculated as:𝐶5,$234,* = 𝐶2J,FGN ∙ 𝑡;G:: + 𝑡KL1Equation 15where Ċoh,fin is the overhead rate for finishing operation (assumed to be the same for thetwo finishing operations).28213 Case study: forming dies for tube bending 3.1	 Classification	of	production	class:	high	mix-low	volume	 The rotary draw bending process is performed on fully electric CNC machines.

Figure 1 showsthe main elements of a typical rotary tube bending machine.Figure 1: Typical rotary tube bending machine (courtesy BLM Group)The straight tube is initially clamped by the bending die (Component 1 in Figure 2) and the clamping die (Component 2 in Figure 2), which draws the tube section according to a predefined radius by rotating. A wiper die (Component 4 in Figure 2) acts as a constrain for the straight part of the tube, allowing a controlled sliding of the tube during the drawing operation. The wiper die fills the gap between the tube and the bending die on the inner radius, to avoid the occurrence of wrinkles.Figure 2:Steps and main components in the rotary tube bending process: 1) Bending die; 2) Clamping die; 3) Pressure die; 4)Wiper dieClamping and bending dies are made of hardened tool steels. Often, the clamping surfaces are coated or case hardened to increase their wear resistance (see Figure 3a and Figure 3b ).

Often, the clamping surfaces are coated or case hardened to increase their wear resistance (see Figure 3a and Figure 3b ). On the other hand, the wiper die is made of bronze alloys (see Figure 3c) or more deformable materials in order to avoid sticking and premature breaks. A starter kit of bending, clamping, and wiper dies is included in each bending machine being sold. Similar kits are provided by machine manufacturers as spare parts and consumables.2822(a)(b)(c)Figure 3: Examples of shapes and material for a) bending dies; b) clamping dies; c) wiper die.Within this work, the entire production in 2015 of these first kits for machines capable to bend average size tubes was analysed. The dimensional range of the considered tubes was between 12.50 mm and 50 mm in diameter and between 15 mm and 250 mm bending radius.

The dimensional range of the considered tubes was between 12.50 mm and 50 mm in diameter and between 15 mm and 250 mm bending radius. Table 2 summarizes the total die production in 2015 in these ranges.Table 2: Total production of starter die kits in 2015 with corresponding number of variants and quantity for each die type.Component Bending dies Clamping dies Wiper dies TotalN of variants 78 130 79 287Total quantity 98 157 105 360Material Tool steel - 40CrMnMo7 Tool steel - 40CrMnMo7 Bronze – Ampco 18Figures in Table 2 puts in evidence that the production of the starter kits is characterised by a large number of variants (in terms of shape and geometry), which are produced in a small batch sizes (in terms of number of parts per year). Moreover, the bending and clamping dies are machined with a 5-axis CNC milling machine from round bars or prismatic blocks of 40CrMnMo7 cold work tool steel in pre-hardened conditions.

Moreover, the bending and clamping dies are machined with a 5-axis CNC milling machine from round bars or prismatic blocks of 40CrMnMo7 cold work tool steel in pre-hardened conditions. After machining the clamping surfaces, these are ion nitrided and eventually mirror-like polished. On the other hand, the wiper dies are obtained through5-axis CNC milling machining of the commercial bronze alloy Ampco 18. Successively, the functional surfaces are mirror-like polished. Both nominal and chemical compositions of 40CrMnMo7 tool steel and Ampco 18 bronze are given in Table 3.Table 3: Nominal chemical composition of the materials originally used (wt. %)Material 40CrMnMo7 Ampco18Si 0.30-0.45 0.20-0.40 1.30-1.60CMnP (max) S (max)0.0350.035CrMo1.80-2.10 0.15-0.25Al10.5CuBal.The analysed industrial case is representative of a high mix-low volume production.

This explain why 5-axis CNC milling machining is the method actually employed to produce all the three classes of dies. Milling allows flexibility in terms of shapes and materials as well as a medium productivity without requiring expensive moulds and dies.3.2	 Families	of	bending,	clamping	and	wiper	dies	 Given the large variability in shapes and material of the investigated components, the bending, clamping and wiper dies were grouped in families similar in geometry, material, manufacturing process and function. The bending dies were grouped into five families (as shown in Figure 4)2823Fe Bal.

The bending dies were grouped into five families (as shown in Figure 4)2823Fe Bal. 3.5whose batch size for the year 2015 is listed in Table 4 along with their weight and height (on the plate).(3) Figure 4 - Families of bending dies: 1) shaped; 2) single-block; 3) multi-block; 4) complete; 5) with inserted tooling(1)(2)(4)(5)Figure 4 and Table 4 evidence similarities in the shape, weight and height for code 1, 2 and 5 while code 3 is significantly more complex in shape, heavier and taller.

On the other hand, code 4 is the smallest, lightest and shortest.Table 4 - Families of bending dies: yearly quantity, weight and maximum height (on the SLM platform)NomenclatureShaped Single-block Multi-block Complete With insert TotalCode1 2 3 4 5Quantity3 8 4 63 20 98Weight [kg] 4.125 3.414 4.593 1.957 3.279Height[mm]124.7 127.1 175.4 110.1 133.1Total weight [kg]12.38 27.31 18.37 123.29 65.58 246.93The clamping dies are classified in two macro-groups: clamping dies to be fixed on the bending die (therefore without a base as support) and clamping dies with a prismatic base as a support.

The clamping die to be fixed on the bending die in Figure 5 is a tool that is inserted on the correspondent bending die (see Figure 4e).Figure 5 – Insert clamping dies for bending diesClamping dies with a prismatic base are characterised by a standard support with a prismatic shape and parametric dimensions, that can be grouped in three classes: small, medium and large (see Figure 6).2824(c) (b) (a) Figure 6 - Families of hybrid clamping dies: a) small; b) medium; c) largeSince the base of the clamping dies in Figure 6 has a prismatic and simple shape, the SLM processing allows for two alternative approaches: 1) fully additive through SLM; 2) hybrid (subtractive + additive) sequence in two steps (as shown in Figure 7): machining of the support (see Figure 7b) through conventional 5 axis CNC milling process and additive manufacturing of the personalized upper part through SLM process directly onto the support surface (as indicate Figure 7c).

To this purpose, an ad hoc platform is required, which houses the machined supports and aligns the parts in the right position for the SLM process.

A preliminary cost analysis, here not reported for the sake of briefness, shown the economic convenience in adopting this hybrid (subtractive + additive) solution with respect to the full additive alternative.Figure 7 – Production method for hybrid clamping dies: a) hybrid clamping dies made up of a bottom support and a toppersonalized part; b) bottom support machined by CNC milling; c) top personalized part manufactured by SLM directly on top of the supportThe clamping dies results in the four families whose batch size for the year 2015 year is listedin Table 5 together with their weight and height.Table 5 - Families of clamping dies: yearly production and weightNomenclatureCode QuantityWeight [kg]Height [mm]Total weight [kg]Insert1430.7388031.73Hybrid-small2180.356306.0Hybrid-medium3580.5343030.96Hybrid-large4380.7823029.72Total15798.82The wiper dies are grouped in three families (as shown in Figure 8) mostly based on the lengthof the wiper die surface.2825(a)(b) Figure 8 - Families of wiper dies: a) small; b) medium; c) large(c)The batch size for each of the three families is similar whereas the long wiper die is significantlyheavier and taller (as listed in Table 6).Table 6 Families of wiper dies: yearly production and weightNomenclatureLongCode1Quantity30Weight [kg] 1.162Height [kg]160Total weight[kg]34.8Intermediate2360.7389826.5Short3390.7299328.4Total10589.93.3	 Year	and	quarter	batch	size	of	the	entire	die	production	 In the year 2015, the production for tube bending dies was planned on a quarterly base.

Therefore, every three months the batch size for the three types of dies was scheduled according to the quantities in Table 7.Table 7 – Quarterly production lot for bending, clamping and wiper diesBending diecode 1 2 3 4 5 Quarter total Year totalQuantity1 2 1 15 5 24 96Clamping diecode 1 2 3 4Quantity11 4 14 1039 156Wiper die code 1 2 3Quantity10 9 827 1083.4	 Number	of	build	job	and	building	time	estimation		 One of the inputs of the cost model is the building time for each build-job tb. This term depends on the division of the yearly production (in this case in quarter lots) and on the available SLM systems. In this study, a Renishaw AM 250 system (Stone, UK) with a 200 W power modulated fiber laser was used. This system is equipped with an inert gas chamber mounting a 250x250 mm2 platform.2826Two commercial, gas atomized powder types were selected with properties similar to the original 40CrMnMo7 tool steel and Ampco18 bronze.

The production of a new powder (in particular for the wiper dies in Ampco18) was not considered because of the high start-up cost of a new powder production. The original 40CrMnMo7 tool steel was substituted by gas atomized 18Ni300 maraging steel powder (Sandvik Osprey, Neath, UK). On the other hand, pure copper powder (LPW, Cheshire, UK) was investigated instead of Ampco18 bronze. Both powders had a particle size between 15 and 45 µm.The chemical composition of both gas-atomised powders are listed in Table 8. Table 9 provides the standard process parameters indicated by the powder supplier in the case of 18Ni300 maraging steel [8]. On the other hand, only indicative process parameters for pure Cu were considered, due to a lack of process data with the given SLM system.Table 8 – Nominal chemical composition of the powders employed in the study (wt.

An optimized mixing of these two categories was obtained by implementing the heuristic nesting of the build preparation software (Magics 20, Materialise, Leuven, Belgium). On the contrary, the hybrid clamping dies and wiper dies were manufactured on separate build-jobs since the first required tailored platforms while the second were built on a different metallic alloy. Therefore, three classes of build-jobs were prepared namely, mixed bending dies with insert, hybrid clamping dies and wiper dies. The final allocation resulted in the total number of build-jobs listed in Table 10.

Total building time and productivity levels are also listed.2827Table 10 Total number of build-jobs per bending dies with insert, hybrid clamping dies and wiper dies in a quarter.Build-job classBending dies with insertNumber of build-jobs8Total volume (cm3) 8976.6Total building-time (h)1173.7Productivity (cm3/h)7.6Hybrid clamping dies42213.4282.47.8Wiper dies23208.1653.64.9Total1414398.12109.7As expected, the SLM processing of maraging steel is more productive than the SLM processing of the copper alloy since the high reflectivity and high thermal diffusivity of copper require slower scanning speed and lower layer thickness.Starting from the allocation of parts indicated in Table 10, the pre-processing, building and post- processing times per each component for the corresponding die families can be evaluated (as shown in Table 11).Table 11: Processing time for the classes of bending, clamping and wiper dies detailing the production planning for a givenquarter.Build-job classBuild-jobVolumeBuilding timePre- process timePost- processVp [mm3]tb [h]tcomp [h]tEDM [h]Bending dies with insert121,698,6621,272,369221.47165.7010.59.71.81.63859,566113.286.71.44859,566113.286.71.45859,566113.286.71.46859,566113.286.71.47859,566113.286.71.481,707,732220.1210.51.4Hybrid clamping dies910642,991703,94181.7389.349.310.20.00.011595,13475.828.60.012271,36035.493.90.0Wiper131,998,364418.538.71.7dies141,209,722235.057.31.3TOTAL14,398,1052109.61112.015.02828Timetmill [h]13.17.98.68.68.68.68.616.00.00.00.00.015.86.7102.4An example of the component layout on the substrate platforms is shown in Figure 9 for eachof the three different classes of platforms.(b) Figure 9 – Allocation on the platform and nesting of the different families: a) mixed disposal for bending dies and insert clamping dies in maraging steel; b) platform with hybrid clamping dies in maraging steel; c) platform with wiper dies in pure copper.(a)(c)Figure 10 shows representative moments of the SLM production: the wiper stand-alone platform, the hybrid plate with support bases ready for the SLM processing and the hybrid clamping dies realised.(b) Figure 10 – Representative moments of the tube bending die production in a mixed and hybrid configuration: a) wiper stand- alone platform, b) hybrid plate with support bases ready to be SLM printed, c) hybrid clamping dies at the end of the SLM process.(a)(c)From the data collected by AddMe Lab, the laboratory for additive processes of the Mechanical Engineering department of the Politecnico di Milano with a Renishaw AM 250 machine indicates that a maximum load factor of 0.77 can be considered for the SLM process .

Therefore, considering a maximum number of hours per year of 8438.4 and a maximum load factor of 0.77, the demand of build jobs in Table 11 is satisfiable by two Renishaw AM 250 systems. A single AM 250 machine with a load factor of 0.96 would be required, whereas with two AM250 the load factor is reduced to 0.48. Therefore, two SLM machines are considered in the present study although their production would be under-loaded.5 Cost model analysisTable 12 includes a detailed list of the main costs and load factors used in the model.

The listedvalues are in line with literature and the data from the industrial partners of the AddMe Lab[9].2829Table 12 – Main costs and load factors used in the modelSymbol Cm Cu Cm 18Ni300 Cre Cplate Cgas cyl Cfilter CC Coh Lre Lplate Lgas cyl Lfilter f rsol Cu rsol 18Ni300 rpow Cu rpow 18Ni300 L twoCost 95 €/kg 69 €/kg 2.50 € 241.50 € 60 € 27 € 500,000 € 8.69 €/h 3 12 4 1 5.4 % 7450 kg/m3 8100 kg/m3 3955 kg/m3 4300 kg/m3 48 % 8 yearsSymbol Ccomp Lcomp two,comp Coh,comp Cmill CEDM Coven Lmill LEDM Coh,finCost 11,600 € 5.1 % 5 years 20.98 €/h 630,000 € 175,000 € 1.46 €/kg 20 % 10 % 20 €The cost analysis applied to the overall annual production allows to conclude that the selective laser melting is more expensive than the conventional milling production process.

In comparison, the SLM additive process results being 60% more costly, as shown in Figure 11.ReferencePre-processingSLMPost-processing180%)%160%(t s o c140%n o i t c u d o r p120%100%la t o t80%e g a t n e c r e P60%40%20%0%SLMReferenceFigure 11 – Total cost of the SLM annual production in respect to the total cost for traditional productionOverall, considering the AM production chain, the cost relative to the SLM process resulted as the higher cost, whereas pre- and post-processing costs were responsible for only 13% of the production total cost (CT,Add, as indicated Figure 12). The contribution given by the four terms to the SLM total cost (CT,SLM) is presented in Figure 13.

The contribution given by the four terms to the SLM total cost (CT,SLM) is presented in Figure 13. The major cost driver consists in the equipment cost (52% of the total SLM cost), while the overhead costs are the second term (30% of the total cost).2830Figure 12 - Division of the production total cost (CT,ADD) in the tree steps: pre-processing, SLM and post-processingSLM	 Overhead 30%SLM	 Material 16%SLM	 Tooling 2%SLM	Equipment 52%Figure 13 - Division of the total SLM production cost (CT,SLM) for cost terms.Contrarily to expectations, the material cost accounts for only 16% of total cost. This is due to the relatively low cost of maraging steel powder, since 77% of the build volume required it (whereas pure copper powder was limited to the remaining 23%).

This is due to the relatively low cost of maraging steel powder, since 77% of the build volume required it (whereas pure copper powder was limited to the remaining 23%). With reference to more demanding fields, such as aerospace and biomedical, where more costly powders are commonly used (such as titanium or nickel-based alloys) in the present study the cost of the powder was not a predominant lever. Finally, the lowest contribution to the total SLM production cost was related to the tooling cost since the cost of goods and materials consumption during build-jobs is almost negligible.Figure 14 shows the cost analysis at build-job level. Orange bars are the reference total production cost in for the milling process. The differences are both related to the SLM productivity and the cost of the same component produced by the conventional method. Indeed, build no 1 and 8 are characterized by similar build times and volumes. Hence, the production2831costs through SLM are similar.

Indeed, build no 1 and 8 are characterized by similar build times and volumes. Hence, the production2831costs through SLM are similar. However, the products obtained in build job 8 cost much less obtained conventionally, resulting in a much higher relative cost.SLM Reference350%300%250%200%150%100%50%0%12345678910111213Build-jobFigure 14 – Cost comparison per build-jobIn order to assess the impact of the most significant inputs of the cost model on the total production cost, a sensitivity analysis was performed. One-factor-at-a-time approach was selected in this initial stage. From the previous analysis, the cost inputs worthy to be explored are the parameters which regulate the equipment C b,SLM,3 and overhead Cb,SLM,4 costs respectively. Hence, the SLM capital cost CC and the load factor L in Equation 8 were varied. The building time present in Equation 8 and Equation 9, was varied as well.

Hence, the SLM capital cost CC and the load factor L in Equation 8 were varied. The building time present in Equation 8 and Equation 9, was varied as well. tb in fact, influences the process productivity once the build volume is assigned (see Table 10). A variation in the material cost was also investigated since two possible scenarios might occur. On one hand, the cost of the powder could increase because a new bronze alloy is developed; on the other hand, the cost of powder could decrease due to an increment in the producers of iron-based powder. Finally, the freedom in design allowed by SLM additive process was investigated through the removal of non-essential material and the use of lattice structures. A lattice structure can be applied to the overall volume of the dies. This results in a reduction in the weight of the components thus reducing material consumption whilst maintaining a high level of mechanical performance and surface finishing.

This results in a reduction in the weight of the components thus reducing material consumption whilst maintaining a high level of mechanical performance and surface finishing. This is particularly beneficial and results in considerably lower material costs and a significant reduction in building time. Table 13 presents the different sources of variability considered in the sensitivity analysis and respective levels.Figure 15 shows the results of the sensitivity analysis. With respect to the baseline total cost of SLM production, the most effective input parameter is productivity. When productivity doubles, the cost of the SLM production cost comes closer to the reference cost (i.e. the actual milling process) but is still not economically viable. The second most influencing factor is the 30% mass reduction which causes a significant reduction on the SLM process cost.

the actual milling process) but is still not economically viable. The second most influencing factor is the 30% mass reduction which causes a significant reduction on the SLM process cost. If we consider a best case variation of the different factors leading to the most favourable possible situation as underlined values in Table 13, the total production cost is reduced to 60% of the reference cost.283214Table 13 – Different sources of uncertainty in the main inputs of the cost model.

Underlined conditions depict the best case.Increment Baseline DecrementSLM system cost500.000 € 400.000 € 300.000 €Load factor 0.80 0.65 0.48Productivity200% 150% 100%Material cost150% 100% 50%Mass reduction0% 15% 30%SLM	MaterialSLM	ToolingSLM	EquipmentSLM	OverheadPreprocessingPostprocessing200%)%(t s o c n o i t c u d o r p150%100%la t o t e g a t n e c r e P50%0%Figure 15 - Variation of different influence factors on the total production cost6 Conclusions In this work, a dedicated cost model for a production route consisting of pre-, SLM, and post-processing stages has been developed. The developed model was applied to a case study consisting in the production of tube bending tools with a large variation in type and small lot size. Different possibilities such as the production of the components without any geometrical modifications, use of hybrid production, and weight reduction were analysed.

Different possibilities such as the production of the components without any geometrical modifications, use of hybrid production, and weight reduction were analysed. The main results can be summarised as follows. • In the case of adopting the SLM process without any geometrical modifications to the parts, the AM process is not economically viable compared to the conventional route based on machining. • The cost decomposition depicts that the SLM processing step is the most influencing factor,which increases the overall part cost.2833The AM process may become economically viable with the current component design. with an increase in the productivity of SLM systems (for example thought the use of multiple lasers) and reduced machine cost.The factor which is most difficult to assess is the added value to the produced parts and therefore stands out as the main challenge for the viability analysis.

Comparative Evaluation of Energy Consumption of Selective Laser Sintering and Injection Molding of Nylon PartsCassandra Tele nko and Carolyn Conner Seepersad Mechanical Engineering Department The University of Texas at Austin Austin, TXAdditive manufacturing to competing manufacturing technologies. This research study focuses on estimating and comparing the energy consumption required for different production volumes of nylon parts using either selective laser sintering (SLS) or injection mo lding (IM) . For IM & SLS , energy consumption is estimated for nylon material refinement and part fabricati on. For fM, e nergy consumption is also estimated for manufacturing the injection mo lds and refining their metal feedstock. A paintball gun handle serves as a representative part for calculating and normalizing material flows and processing times.

A paintball gun handle serves as a representative part for calculating and normalizing material flows and processing times. For different sets of assumptions, cross-over production vol umes are calcu lated, at which the per-part e nergy consumption of the two processes is equivalent. These energy-based cross-over production vo lumes are compared to simila r economic cross-over producti on volumes available in the literature.is often advocated as a sustainable alternative1. Introduction$elective laser si ntering (SLS) is a prominent techno logy for additive manufacturing (AM) of functional parts. SLS and competing AM technologies are generally assumed to be more environme ntally sustainable than conventional manufacturing methods because the additive process minimizes tooling, material waste, and chemical fluids.

Quantitative support fo r many of these hypotheses is not publicly avai lable (Drizo and Pegna, 2006); accordingly, a 2009 NSF-sponsored workshop ide ntified multiple research needs relating to sustainability, including material performance data, measures of process sustainability, and comparisons with other manufacturing methods (Bourell et al. , 2009). The research presented in this paper addresses some of these c hallenges by evaluating the energy consumption required to fabricate nylon parts using SLS and comparing it with that required for inj ection mo lding (IM) the same parts.Estimates of energy consumption are obtained from life cycle inventories (LCis) of SLS and IM. An LC I is a n important part of life cycle analysis (LCA), an accepted method for quantifying the e nvironmenta l impacts o f a product o r process througho ut its life cycle (ISO, 1997).

An LC I is a n important part of life cycle analysis (LCA), an accepted method for quantifying the e nvironmenta l impacts o f a product o r process througho ut its life cycle (ISO, 1997). A traditional product life cyc le, shown in Figure 1, starts with procurement of materials from the earth and ends with return of materials to the environment or re-processing plant. In the context of the life cycle, an LCl e ntails tracking the flows of e nergy and/o r materials between a technical system and its surroundings. In this research , LCI data are collected and evaluated for energy consumption during two stages of the SLS a nd IM life cycles: material refinement and part fabrication.

In this research , LCI data are collected and evaluated for energy consumption during two stages of the SLS a nd IM life cycles: material refinement and part fabrication. Specifically, energy consumption is estimated for refinement of nylon feedstock material, use of SLS and IM equipme nt for part fabricati on, and fabrication of an injection mo ld for a representative part.Materials ItOutputsAir EmissionsSoil EmissionsWater Emissio nsSolid WasteFigure I: Inputs, outputs and processes included in a product's life cycleA comparative study of SLS and IM is interesting because several differences between SLS and IM manufacturing systems affect their energy consumption. An SLS machine consumes a significant amount of energy while fabricating parts, but SLS does not require tooling. IM machines consume less energy while fabricating parts, but 1M requires metal molds.

An SLS machine consumes a significant amount of energy while fabricating parts, but SLS does not require tooling. IM machines consume less energy while fabricating parts, but 1M requires metal molds. These metal molds require significant amounts of energy investment as a result of the metal feedstock and machining operations used to fabricate the mold. For both JM and SLS, the energy consumed per part depends on the number of parts fabricated. IM requires large numbers of parts to justify the investment in the mold. SLS energy consumption per part depends on the density of its build and related factors such as part orientation. All of these factors are considered in this study.Previous work has included quantification of the material and energy use of laser sintering (Luo et al. , 1999; Mognol et al.

All of these factors are considered in this study.Previous work has included quantification of the material and energy use of laser sintering (Luo et al. , 1999; Mognol et al. , 2006; Kellens eta/., 2010; Sreenivasan and Bourell, 2009; Baumers et al., 20 1 0; Dotchev and Yusssof, 2009), IM (Thiriez and Gutowski, 2006), and die manufacture (Dalquist and Gutowski, 2004; Morrow et al., 2007), but SLS and IM have not been systematically compared in an LCI. Hopkinson and Dickens (2003; 2006) and Atzeni et al. (2010) compared the monetary costs of SLS and IM, but not the energy costs. Morrow et al. (2007) compared the energy consumption for machining and Direct Metal Deposition of a sim ple steel die, but they did not consider the significant material refinement costs of metals (Dahmus and Gutowski, 2004) nor did they consider IM or SLS processes.

The goal of this paper is to build from these findings and create a comparative study of IM and SLS energy requirements for material refinement, tooling, and nylon part production. The scope of the study is defined in Section 2, followed by the detail s of the LCI in Section 3 and the results in Section 4.2. Goal and Scope of the Life Cycle Inventory (LCI)The goal of this study is to understand whether SLS is more energy efficient than IM, and if so, under what circumstances and for what production volumes. The scope of this study is limited to material refinement and part fabrication stages of the life cycle of a functional nylon part. Specifically, the scope is limited to evaluating the energy consumed for refining nylonfeedstock, fabricating parts with SLS and IM equipment, and fabricating injection mo lds for IM. T hese processes and corresponding reference fl ows are depicted in Figures 2 and 3.

T hese processes and corresponding reference fl ows are depicted in Figures 2 and 3. This energy study does not inc lude packaging, storage, transportation, or use of the functional part; nor does it include recycling of the functional part, treatment of waste streams from the material refinement or part fabrication stages, or the energy consumption of the infrastructure (e.g., climate control) surrounding the SLS o r TM machines. Parts are assumed to be fabricated from nylon on a 3DSystems® Sinterstation® HiQ™ + HiS™ SLS machine.Figure 2: The scope and reference flows of the SLS LCI,.-----------1I__ 7Final Part (35g)Material WasteFigure 3: The scope and reference flows fo r the IM LCIA representative part was selected as the basis for the functional unit of the LCI. As shown in Figure 4, the representative part is a paintball gun hand le created by undergraduate and graduate students in a Solid Freeform Fabrication course taught by one of the authors (Bazan et al.

, 2009). As shown in Figure 4, the part is created in two halves and encloses the metal frame of a paintball gun handle. This part was se lected for its moderate size and complexity and suitability for additive manufacturing. For example, using 3D scans of molds of customers' hands, form-fitti ng grips could be added to the hand les, along with personalized names or insignia. In this study, however, the part is assumed to be mass-produced so that only one injection mold must be designed and fabricated.The functional unit used in the LCI is the number of representative parts manufactured, also referred to as the production volume. The handles are small enough, approx imately 3.2cm by 2.54cm by 12.7cm per half, that a single 3DSystems® Sinterstation® HiQ™ + HiS™ SLS machine could produce 300 halves or 150 units within t he machine's bu ild volume of 38 1mm(W) by 330 mm (D) by 457mm (H). This estimate allows fo r l2.7 mm of spacing around the edges of the bui ld chamber.

This estimate allows fo r l2.7 mm of spacing around the edges of the bui ld chamber. The spacing between halves is assumed to be 2.54 mm on all sides, with each half oriented with its longest dimension parallel to the height axis of the buildchamber to increase packing density. In this configuration, 50 un its (or 100 halves) can be fabricated in a single row, approximately 130 mm high, in the SLS build chamber, w ith 3 rows constituting a full build of 150 units. A small or partial build is defined as a single row of 50 units.Figure 4: The paintball gun handle consists of two hal ves per part3.

A small or partial build is defined as a single row of 50 units.Figure 4: The paintball gun handle consists of two hal ves per part3. Life Cycle Inventory MethodologyBased on the assumptions described in Section 2, an LCI is performed for SLS and JM.The total energy consumed by each process is quantified by Equations 1 and 2:EAM = Enylon + EsLs ElM = Emetal + Emachining + Enylon + EM(l) (2)where E AM denotes the total energy required to create SLS parts; Enylon denotes the energy consumed during the processing of nylon feedstock for the final part; EsLS denotes the energy consumed by the SLS unit manufacturing process; E 1Mdenotes the total energy required to create IM parts; Emera/ denotes the e nergy consumed during the process ing of steel or alum inum feedstock for the injection molds; Emachiningdenotes the energy consumed during machining of the injection molds; and EM denotes the energy consumed by the IM unit manufacturing process.

The following subsections describe the process for calculating the input parameters for Equations I and 2.3.1 Energy Consumption of Nylon ProductionBoth processes are assumed to use the same Nylo n 12 material and refining processes. These processes are surrounded by dashed lines in Figures 2 and 3. Equation 3 is used to determine the energy consumption of material refinement :Enylon = m feedstock (Y1.nylon_productwn) (knylon )(3)where mfeedsrock is the mass of nylon required by the IM or SLS manufacturing process; Ynylon_producrion is the material yield of nylon production measured as the mass fraction of nylon feedstock that is successfully converted to nylon powder or granulate; and k,ylon represents the specific energy consumption (SEC) of the nylon production process.The SEC is the amount of energy, in megaj ou les, consumed during the processing of a single kilogram of nylon feedstock.

The Plastics Europe life-cycle inventory (GaBi, 2006; Hischier, 2007) for nylon 6 granulate production provided an SEC for nylon production, knylomof 116 MJ per kilogram of nylon produced. The difference in SEC between Nylon 6 and Nylo n 12 was assumed to be negligible.The nylon production yield, Ynylon__producrion, was estimated separately for IM and SLS. Material loss during nylon production differs for nylon granulate and nylon powder, utilized by IM and SLS, respectively. Since the estimate for k nylon incorporates the yield of the granulate production process, the material yield of IM nylon granul ate production, YnylonyroducrianOM), was assumed to be I 00%. In contrast, the material yield for SLS nylon powder production, Ynylon_proauction(SLS), was assumed to be 98%. Th is assumptio n was based o the details of U.S. Patent 4,334,056, wh ich describes the process of creating nylon powder for SLS (Meyer et al., 1982; Zarringhalam et al. 2006).

Th is assumptio n was based o the details of U.S. Patent 4,334,056, wh ich describes the process of creating nylon powder for SLS (Meyer et al., 1982; Zarringhalam et al. 2006). The patent suggests that 2% of particles yielded by th is process are outside of the acceptab le size range for SLS.The total amount of feedstock is calculated acco rding to Equation 4:mfeedstock = nparts ffipart (y1) SLSor/M(4)where nparrs is the number of parts; mpart is the mass of each part (35 g for the representative part); and YsLSoriM is the process material yield for either SLS or IM.A yield rate, Y1M, of 90% is assumed for IM . Waste varies by shop and is created during machine start-up, purging, and maintenance. O lmstead and Davis (200 I) cite a "standard reject rate" of 5%. Michaeli and Greif (2001) cite that waste content of material can be 5-50%, presumably by weight.

O lmstead and Davis (200 I) cite a "standard reject rate" of 5%. Michaeli and Greif (2001) cite that waste content of material can be 5-50%, presumably by weight. A 90% yield rate is based on the work of Th iriez and Gutowski (2006), who perfo rmed an energy LCA of JM using data from over 1 00 sources and estimated a scrap rate of I 0% for IM manufacture.The yield rate of SLS is estimated fro m best practice and powder utilization data. Best practice requires that the powder for each build contain at least 30% virgin material, assuming that the powder is not infinitely recyclable. Furthermore, depending on the dens ity of the build, Dotchev and Yussof (2009) re port that as little as 12.8% of the powder is sintered into fina l parts, indicating that some of the used powder is eventually discarded to accommodate the virgi n material requirement and the powder losses experi enced during part breakout.

Based on powder utilization data reported by Dotchev and Yussof (2009) and the 30% virgi n material requirement, the authors calculated that between 40% and 80% of the material in a single build is sintered into a final part in the initial build or subsequent builds. In Equation 4, a material yield, YsLS, of 60% (40% material loss) is assumed, wh ich agrees closely with values suggested by a local service bureau.3. 2 Energy Consumption of SLSAs reported in Equation I, the energy required to fabricate an SLS part, EAM, is the sum of the energy embedded in the nylon feedstock, Enylon, and the energy consumed during the SLS process, Esls·It is difficult to prescribe a specific energy consumption value fo r SLS in terms of part volume or mass, because energy consumption varies w ith build density and height. For example, Mognol eta/.

For example, Mognol eta/. (2006) built the same metal sintered part in multiple orientations and found that the difference in build height corresponded to a range of 115-187 MJ of energy consumption for the build . Orientation alone accounted for a 60% increase in energy use, because changes in orientation can cause changes in the height and duration of the bui ld and therefore increase the energy consumption of the layer-based SLS process.Accord ing ly, the energy consumption of the SLS unit process, EsLS, was calculated by summing the prod uct of average pre-heating powe r consumpti on and preheat ti me w ith the product of the build time and average build power consumption as shown in Equation 5:EsLS = t buildEbuild + t preheat Epreheat(5)where E build is the average power required during the build stage; t build is the bu ild time; Epreheat is the average power required to preheat the build chamber; and !preheat is the time requ ired to preheat the build.

Equation 5 is further refined into Equation 6:EsLS = nzayers tlayer E build+ tpreheatEpreheat(6)where the time to build is calculated as the time to build each layer, l taym mu ltiplied by the number of layers in a bu ild, n tayers·For this study, average power consumption and build and preheat times were measured from three builds on a 30Systems® Sinterstation® HiQ™ + HiS™ SLS machine. A F luke I 75 0 3-phase power meter was used to measure the voltage and current drawn by the mach ine during two builds and by the laser chiller during a third build. As re ported in Table 1, bui ld densities and heights ranged from 9% to 13% and 60 to 152 mm, respectivel y. The warm up time for these builds ranged from 2 ho urs to 2 hours a nd 45 minutes. Layer scan and preparation time required an average of 45 seconds during the first two builds. The machine was found to draw an average of 4 kW during the warm up stage and 3.5 kW during the build stage.

The machine was found to draw an average of 4 kW during the warm up stage and 3.5 kW during the build stage. The chiller was found to operate at a nearly constant 2 kW for both stages. During the third build of a pyramid like structure of layers ranging from 90-20% density, layer preparation a nd scan time was fo und to range from 60-25 seconds as density of the scan area decreased .Table I: Results for power and chi ller measurements using a Fluke 1750, 3 phase power meter and a 3DSystems® Sinterstatio n® HiQTM + H iSTM SLS machineWarm Up StageBuild StageBuild Chiller Duration Power Duration Power Packing Density Build Height Avg Layer Time1 2 32 hours 3.9 kW13h3.5 kW9%152 mm46 s····· ........... ·2ti .4siii·· ·ia·kw ··a11·1·2r-n·· ·3:5 !(vii ·· · ··· · ··· 13o/~ ·········· ·····9i r-n·n;· ···· ·········· 45·5··········· .. ... ~. -~ ... ... ........ .... ...... . ...... ............... . ,, ... ......... ... ... ........... ....... ..... ............ ......

... ~. -~ ... ... ........ .... ...... . ...... ............... . ,, ... ......... ... ... ........... ....... ..... ............ ...... ... .. .... ... .... .2 kW2 hours3h 45m13%35 mm56 sFrom these experimental observations, the machine a nd chiller were assumed to operate at a combined average of 6 kilowatts while preheating the bui ld, Epreheat , and 5.5 kilowatts during the build, Ebuild· Each bui ld was assumed to incur a warm up time of two hours, lpreheac, and each layer was assumed to require 45 s for preparation and scanning, t 1ayer · As explained in Section 2, the build chamber accommodated 150 re presentative parts, arranged in 3 rows of 50 parts per row. Each part row was approximately 130 mm high a nd included 860 layers, n tayers. each of which required approximately 45 seconds to build and prepare.

Each part row was approximately 130 mm high a nd included 860 layers, n tayers. each of which required approximately 45 seconds to build and prepare. Energy was calculated for two scenarios, partial builds of a sing le row of 50 parts and fu ll builds of three rows of parts, for a total of 150 parts.The experimentally de termined values of power and energy consumption for the 30 Systems® Sinterstation® HiQ™ + Hi S™ SLS machine are similar to those published in the literature for compara ble appl ications. Prev ious studies (Luo eta/. , 1999; Baume rs et al., 2010; Mognol et al. , 2006 ; Kellens et af., 20 I 0; Sreenivasa n and Bourell , 2009) of metal and plastic sintering have measured power draws rang ing from 3 to 19 kW during the build stages of plastic and metal sintering machines including the EOS EOSlNT P760, EOS INT M250 Xtend, 3DSystems® Sinterstation® HiQ™ + HiS™ SLS machine, and DTM Sinterstation 2500. Luo etal. ( 1999) and Kellens et al.

Luo etal. ( 1999) and Kellens et al. (20 I 0) fabricated polymer parts on a Sinterstation 2000 and an EOSINT P760, respectively, and repo rt SLS machine energy consumption that can be expressed as a specific energy consumption value rang ing from I 08 to 144 MJ per kilogram of fabricated part. Based on Equation 6 and the assumptions reported in this section, the specific energy consumpt ion of the 3DSystems® Sinterstation® HiQ™ + HiS™ SLS system is approx imately 130 MJ/kg for fabrication of the representative part.3.3 Energy Consumption of JM Mold ProductionAs reported in Equation 2, the total energy required to make an lM part, ErM, includes the energy consumed to produce the mo ld (Emetat+ Emachining) , the energy required to refine the material, Enylon, and the energy consumed by the IM process itself, E M. Calculation of energy consumption for nylon refinement, E nyt011 , was described in Section 3. 1. To calculate Emetal and Emachining.

Calculation of energy consumption for nylon refinement, E nyt011 , was described in Section 3. 1. To calculate Emetal and Emachining. the two-plate injection mold for this study is assumed to fit two copies (four halves) of the paintball handle. Equation 7 was used to calculate Emewt, the energy required to create the necessary vo lume of tool steel entering the machining process:Emetal = m metal ( (1- r)kvirgin + rkrecyc )(7)where m111erat is the mass of metal required for the injection molds before machining; r is the fraction. of recycled co ntent of each metal; and kvirgin and krecyc denote the specific energy consumptio n (S EC) or megajoules of energy required to refine a ki logram of metal from virgin and recycled sources, respectively.The required mass of metal was calculated using dimensio ning and tolerancing guidel ines from the text by Kazmer (2007).

Each mold plate was assumed to contain cavities or cores for two complete re presentative parts (i.e., four halves of paintball gun handles). The dimensio ns of each mold plate allowed for perimeter spacing equal to the depth of the caviti es and minimal d iameter cooling channels. The total vo lume of steel or aluminum for each plate was estimated to be approximately 4,000 cm3 , resul ting in approximate masses of 3 1 kg for steel or 11 kg for aluminum.T he SECs for stee l and aluminum productio n are cited from Dahmus and Gutowski ' s (2004) study of machining. The SEC of a metal varies significantly with recyc led content. For example, virgin aluminum embodies approximately 270 megajo ules per kilogram, kvirg;11 , and recycled aluminum embodies approximately 16 MJ/kg, krecyc· The recycled contents for steel and aluminum can be as high as 80% and 20%, respectively.

For stee l the recycled material requires 9 MJ/kg, krecyc, and the virg in material requires 31 MJ/kg, kvirgin· Three possib le materials are considered for the mo lds: aluminum with 20% recycled content, steel w ith 80% recycled content, and I 00% virgin tool steel.The energy for mach ining the injection mo lds, Emachining. is also evaluated using values from Dahmus and Gutowski (2004). The total machining energy is then ca lculated using Equation 8:Emachining = k machining ( V cavity + Vcore)(8)where Vcavity denotes the volume of material removed to create the part cavities in the mo ld; Vcore denotes the volume of material removed to create the cores in the mold; and kmachining denotes the SEC of the machining process.To manufacture two paintball gun handles in a sing le mo ld, 270 cm3 wou ld be removed to create the four cavities, and 2,500 cm 3 wo uld be removed to create the four cores. Manufacture of the cooling channels and runner system are not considered.

Manufacture of the cooling channels and runner system are not considered. The SEC for machining, kmachining.changes with the hardness of the material being machined. Steel requires more machining energy than aluminum. Dahmus and Gutowski (2004) report four average SEC values for machining. The modal values for aluminum, 5 kJ/cm 3 IM Process-Re lated Energy Consumption, and steel, 20 kJ/cm3, are used in this study.The ene rgy consumed by the IM machine itse lf is ca lculated as fo llows:EM = mreedstockk/M(9)where kiM is the SEC of the IM manufacturing process. T hiriez and Gutowksi (2006) repo rt a wide range of SEC values for IM. They identify three types of IM machines in use in the Un ited States: electric, hydraulic, and hybrid. They report average, high and low values of SEC for each machine type, and they assume that 70% of the machines in use a re hydrau lic. Therefore, the average SEC value, ll MJ/kg, for hydraulic IM machines is used in Equation 9.4.

Therefore, the average SEC value, ll MJ/kg, for hydraulic IM machines is used in Equation 9.4. Energy Comparison of SLS and IMEquations 1 through 9 are used to calculate the energy required to fabricate different quantities of representative parts w ith SLS or IM. Figures 5 and 6 illustrate total energy consumption, excluding and includ ing mold production, respectively. Each column in the figures represents the energy calculated us ing Equations I a nd 2 for SLS and IM, respectively. These columns are decomposed into "Nylon Production" from Equation 3 for both SLS and IM, "Build Preheat" from the second term of Equation 6 for SLS, and "Part Manufacture" from the first term of Equation 6 for SLS and from Equation 9 for IM.

Ene rgy consumption is calculated for two different production volumes of representative parts: either a partial/small SLS build of 50 parts or a fu ll SLS build of 150 parts, as defined in Section 2.200018001600U) Q)14001200·-ns~ 01000C) Q) ~ 600800400 t ... .200 i0185 MJD1M Small Build555 MJ DIM Full Build602MJSLS Small BuildSLS Full BuildPart ManufactureBu ild Preheat20M I59 MJ214 MJ43 MJ641 MJ 43 MJDNylon Production165 MJ496MJ345 MJ1036MJFigu re 5: The energy breakdown for IM and SLS of small (nparts =50) and full (nparts =150) builds of a representative part. The IM energy totals exclude mold production.As shown in Figure 5, SLS consumes more e nergy per part than IM whe n mo ld fabrication is not cons idered. A full or partial build in SLS consumes more than three times the energy required to IM the same number of pa rts.

A full or partial build in SLS consumes more than three times the energy required to IM the same number of pa rts. Material inefficiency resu lts in SLS requiring approximate ly twice the nylon production of IM ; furthermore, the part manufacturing process requires more than I 0 times the energy of IM manufacturing. Taller builds of SLS are marginally more efficient per part because build preheat energy requirements are assumed to be equivalent for partial and full builds. It is important to remember that both full and partial builds are packed densely in the build chamber, as described in Section 2, and that less dense SLS builds would consume more energy per part, w ith energy consumption proportional to the height of the build during the build stage.When injection mold fabrication is included in the energy consumption calculations, SLS consumes less energy than IM for small production volumes.

. ~-. ~- .~. - ~~ • • A641 MJ 43 MJ 1036 MJFigure 6: The energy breakdown for IM and SLS of small (nJXlrls =50) and fu ll (n"""·' = 150) builds of a representative part. The IM energy totals include mold production.Injection mold fabrication requires significant energy consumption when compared to bu ilds of SLS parts. T he 80% recycled steel injection mold shown in Figure 6 is the least energy intensive of the three mold types considered, but it alone requires approx imately 460 MJ of energy or 75% of the total e nergy required for a small SLS build of 50 representative parts, including SLS part manufacture and nylon productio n. Mold fabrication from virg in steel or 20% recycledaluminum requires approximately 2000 or 2300 MJ of energy, respectively- more than the total energy consumed for a full SLS build of 150 representative parts.

When the energy consumption of nylon production and part manufacture are added to that of mold fabrication, the tradeoffs favor SLS even more.Although SLS uses significantly more energy than IM during part manufacture, the initial energy investment for IM mold plate manufacture creates a fixed energy investment that is not present in SLS. This fixed energy investment allows SLS to be more energy efficient per part for small production volumes. Crossover production volumes, for which TM and SLS use equivalent amounts of energy, are illustrated in Figure 7. The horizontal markings in Figure 7 highlight the initial energy investment required to fabricate the mold, and show that a significant number of SLS parts could be manufactured with less total energy than that required to fabricate the mold alone for IM. Depending upon the metal used for the mold, SLS uses less energy than TM when production volumes can be serviced by only a few builds.

Depending upon the metal used for the mold, SLS uses less energy than TM when production volumes can be serviced by only a few builds. For the representative part, the energy consumption per part for the aluminum injection mold scenario is similar to the SLS energy consumption per part for production volumes of approximately 300 parts. The virgin steel injection mold scenario is only slightly less energy intensive and incurs similar crossover production volumes. In contrast, the recycled steel injection mold substantially reduces the crossover production volume to approximately 50 parts. These results indicate that only products with very small production volumes can be manufactured more energy efficiently using SLS.o Rec. Steel 1M6000Recyc. AI. IM... • • ++ Virgin Steel 1M5000A SLS (Small Build)SLS (Full Builds)4000+++~3000+ .. +--- ...... _+_ .. - - - • .4 -- .. -- - .. ... ... .... .. .... .. ..Recyc. AI. Mold. .. .. - ~. -;; ..... ~-- ~6.

1 Atzeni et al. (20 1 0), Ruffo et al. (2006), and Hopkinson and Dickens (2003; 2006) observed similar trends in monetary cost comparisons of SLS and IM. However, the magnitudes of the crossover production volumes differ significantly between those monetary cost studies and the energy study reported in this paper. Hopkinson and Dickens (2003; 2006) calculated machine, labor, and material costs for IM and laser sintering of a small part with maximum bounding dimension of 35 mm. They calculated a monetary crossover volume of approximately 14,000 parts. Monetary crossover volumes dropped by more than 80% for larger parts ( 44g and 210 mm maximum bounding dimension), more equivalent to the geometry of the representative part in this study. Ruffo et al. (2006) modified the Hopkinson and Dickens study to account for powder recycling and other costs and to adjust costs for low production volumes, less than a full build.

Ruffo et al. (2006) modified the Hopkinson and Dickens study to account for powder recycling and other costs and to adjust costs for low production volumes, less than a full build. They calculated crossover volumes as low as approximately 9,000 parts for the 35 mm part. Atzeni et al. (2010) redesigned a multi-part assembly to reduce part count in SLS and added assembly costs to their monetary cost comparison of SLS and IM. With assembly costs, crossover volumes reached as high as 60,000 or more parts (Atzeni et al., 2010). Although these studies differ in scope, the large discrepancy between monetary and energy crossover volumes indicates that SLS may be more cost effective than energy efficient within the boundaries of this study.50 .. ! 40 t - 35 .I I II.. ta 30 - c.. ~ ........... 25 ~ 20 ] l55 ''l '0 Rec. Steel I MRecyc. AI. I M45 .I+ Virgin Steel 1Mt~~- SLS (Small Build)--- • SLS (Full Builds) •• • • a.wj~ ,, I+Ill • ,. ~ •10 b •• .........

A smaller, less complex part would decrease the energy required to fabricate the mold for IM, thereby tending to decrease the crossover production volume. In contrast, SLS energy consumption per part is strongly dependent on the number of parts that can be packed into a single build, with smaller1 SLS energy consumption per part is relatively constant when one assumes that SLS builds are always equivalently dense.

In this study, production volumes are increased in increments of small builds (i.e., a fully dense row of 50 parts in a partial build) or full builds (i.e., a full build of 150 parts at maximum practical packing density).parts typically utilizing the build volume most efficiently, tending to increase the crossover production volume.As a point of comparison, crossover production volumes we~e calculated for a second representative part, the USB drive cover illustrated in Figure 9, which is smaller than the paintball handle used for the previous analysis.Figure 9: USB case used for comparison (Blogspot, 2011).The bounding dimensions of the USB drive cover are 4 mm by 16 mm by 45 mm, and it weighs approximately 1 g. At maximum packing density, 7524 parts can be fabricated in a single build on the 3DSystems® Sinterstation® HiQ™ + HiS™ SLS machine, with the parts arranged in 9 If partial SLS builds are densely packed rows, each of which constitutes a partial build.

assumed, along with injection mold cores and cavities that can accommodate 20 USB covers, crossover volumes range from approximately 1500 to 3200 parts for virgin steel and 20% recycled aluminum injection molds, respectively. These values are much larger than the crossover production volumes for the paintball gun handle because the USB covers are much smaller and can be fabricated in larger numbers in a single SLS build, Additionally, the shallow size of the USB cover allows for thinner injection mold plates and a smaller energy investment for IM.5. ClosureThe results of this comparative LCI of SLS and IM indicate that manufacturers can save energy using SLS for parts with small production volumes. For the representative part in this study, the crossover production volume at which SLS and 1M consumed equivalent amounts of energy ranged from approximately one to two full builds, or 150 to 300 parts, depending upon the material used to fabricate the injection mold.

Energy crossover production volumes are much larger for a smaller representative part, indicating that specific crossover production volumes are also sensitive to the size and geometry of the representative part. In both cases, the energy crossover production volumes are much smaller than economic crossover production volumes for SLS and IM, indicating that SLS may be more efficient from a cost perspective than an energy perspective, relative to IM. The large material waste, high power draws, and long operating times of SLS make it generally inefficient to use SLS for large production volumes.The results and analysis support three major recommendations for reducing the energy consumption of SLS. First, build volumes should be packed as densely as possible to maximize the part output per build height because SLS energy consumption is dependent upon the height of the build and the corresponding number of SLS layers.

Second, material-related energy consumption could be improved by engineering infinitely recyclable powder and reducing powder loss during handling. Infinitely recyclable powder could reduce powder scrap rates from 40% to 10%, the amount of powder reported to be lost during part break-out. Third, reducing the time required to scan and prepare each layer would significantly lower the energy consumptionof SLS. A lthough material efficiency can save money and energy, the effect of material loss is less significant than build time.Future work is needed to refine this energy analysis and obtain a broader understanding of the relationship between part design and energy use. Studies that consider mixed part bui lds, a finer resolution of production vo lumes, and parts and builds from current manufacturing centers could provide further insight into SLS energy use.

Studies that consider mixed part bui lds, a finer resolution of production vo lumes, and parts and builds from current manufacturing centers could provide further insight into SLS energy use. The scope of future energy studies cou ld also be extended to include factors such as di stribution, waste hand ling, and infrastructure that were not considered in this study.Acknowledgments Cassandra Telenko is supported by an NSF Graduate Research Fellowship.ReferencesAtzeni , E., L. Iuliano, P. Minetola, and A. Salmi , 20 I 0, "Redesign and Cost Estimation of Rapid Manufactured Plastic Parts", Rapid Prototyping Journal, Vo l. I 6 fssue 5 pp. 308-317Baumers, M., C. Tuck, R. Hague, I. Ashcroft, and R. Wi ldman, 20 10, "A Comparative Study of Metallic Additive Manufacturing Power Consumption," Proceedings of the 2010 Solid Freeform Fabrication Symposium , August 9-11 , The University of Texas at Austin, Austin, TX.Bazan, F., F. Concha, Z. Silva, M. VanOverloop, and R.

423-428 .Kirk, R., and D. Othmer, 2007, " Polyam ides, P lastics" Kirk-Othmer Encyclopedia of ChemicalTechnology, Vo lume 19, John Wiley & Sons, Inc, USA.Luo, Y., Z. Ji, L. M ing, and R. Caud ill, 1999, "Env ironmental Performance Analys is of So lid Freeform Fabrication Processes," P roceedings of the 1999 IEEE International Symposium on Electronics and the Environment, May 11-1 3.Meyer, K., K. Hornung, R. Feldmann, and H. Smigerski, "Method for Polytrop ically Precipitating Polyamide Powder Coating Compositions Where the Polyamides Have at Least I 0 Aliphatically Bound Carbon Atoms Per Carbonamide Group," US Patent 4,334,056.M ichae li, W ., and H. Greif, 200 I, Training in Injection Molding: A Text and Workbook, HanserGardner Publications, Muni ch.Mognol, P., D. Lepicart, a nd N. Perry, 2006, "Rap id P rototyping: Energy and Environment in theSpotlight," Rapid Prototyping Journal, Vol. 12, No. I, pp. 26-34.Morrow, W., H. Ki m, 1. Mazumder, and S.

COMPARATIVE FINITE ELEMENT STUDY OF CUBIC UNIT CELLS FOR SELECTIVE LASER MELTINGA.O., Aremu, I. Maskery, C. Tuck, I. A., Ashcroft, R.D., Wildman, and R.I.M, HagueEPSRC Centre for Innovative Manufacturing in Additive Manufacturing, Faculty of Engineering, University of Nottingham, Nottingham, NG7 2RD, UK.AbstractSelective laser melting (SLM) enables the utilization of complicated lattice structures in metallic components. To exploit this capability, it is important to understand the struc- tural properties of these lattices. Topological variations in lattices are diverse, however, only a few are suitable for SLM since some lattices require supports during manufacture while others self-support. Diﬃculties associated with the removal of these supports and their detrimental eﬀects on surface ﬁnish makes the latter group better suited for SLM.

Diﬃculties associated with the removal of these supports and their detrimental eﬀects on surface ﬁnish makes the latter group better suited for SLM. In this work, we investigate the structural properties of some self-supporting unit cells via a ﬁnite element study and show that the performance of a lattice structure is largely dependent on the topology of the unit cell. Variants of the gyroid and face centred cubic unit cells performed better than body centred cubic cells. This was also observed when lattices, made of repeating unit cells were compared.1IntroductionAdditive manufacturing (AM) techniques for metals could improve the mechanical performance of components, since they are less reliant on constraints imposed by tradi- tional techniques. However, ideas are needed to overcome barriers encountered with AM for metals.

However, ideas are needed to overcome barriers encountered with AM for metals. The need for supports during manufacture is a prevalent issue as this has a detrimental eﬀect on the surface ﬁnish [1]; and in some cases, where the topology of the component is highly complex, such supports are diﬃcult to remove. Support structures are used to preserve the structural integrity of a component experiencing thermal stresses during manufacture [2]. By suitably orientating the component, the problem can be mit- igated, yet supports are required to separate the component from the build platform. Restricting supports to accessible regions of a component’s domain is beneﬁcial. This is particularly useful when selective laser melting (SLM) is used to produce components embedded with lattice structures, removing supports from lattices is impractical.There is a growing interest in understanding the structural properties of lattices pro- duced with SLM.

The build chamber, laser settings and morphology of the powder could aﬀect the quality of the lattices, however, it is the choice of material, topology and relative density of the lattice that greatly inﬂuence structural properties [3]. Materials develop- ment for SLM is an active area of research with stainless steel, titanium and nickel alloys being used for lattices as seen in [4, 5, 6, 7]. Hasan et al. [6] observed that heat treating a body centered cubic (BCC) lattice composed of a titanium alloy improved it microstruc- tural properties while dispersing contaminates. Gumrik and Mines [7] studied the BCC1238lattice and proposed a theoretical model for the compressive behaviour of the lattice with stainless steel as the base material. Stainless steel was also employed by Yan and Hao [8] to understand the dependence of the structural properties of gyroid lattices on the cell size.

Stainless steel was also employed by Yan and Hao [8] to understand the dependence of the structural properties of gyroid lattices on the cell size. The eﬀective yield strength and modulus of the gyroid were observed to decrease as the cell size increased. Smith and Guan [9] studied the BCC lattice and its variant, the BCCz, and observed that the stiﬀness and yield strength of the lattices signiﬁcantly improved when the aspect ratio of the unit cell was reduced.Gyroid and BCC lattices and some of their variants require no supports and have therefore been studied extensively. However, there is a broader class of self-supporting lattices with potentially better structural properties. Also, Thomas [2] outlined design rules to achieve self-supporting designs. This creates an opportunity to develop lattices for SLM, yet it is important to understand the structural beneﬁts of existing lattices.

This creates an opportunity to develop lattices for SLM, yet it is important to understand the structural beneﬁts of existing lattices. In this paper, the stiﬀness of the BCC, face centered cubic (F CC) and gyroid lattices are compared via ﬁnite element analysis.2 Self-supporting Unit cellsTriangilar facets were generated to deﬁne the topology of seven self-supporting cubic unit cells. The length, width and height of cell was set to 10 mm. Facets were subse- quently written into STL ﬁles with relative density, 0.23, to ensure a fair comparison can be made across the cells. The relative density is the ratio of the density of the cell to that of a solid cube with same dimensions. The structural stiﬀness of seven self-supporting unit cells was then determined and compared. Cells include variants of BCC, F CC and gyroid lattices as seen in Fig. 1. The BCC cell is composed of eight cylindrical struts which connect the center of a cube to its corners and trimmed to ﬁt the cube (Fig.

1. The BCC cell is composed of eight cylindrical struts which connect the center of a cube to its corners and trimmed to ﬁt the cube (Fig. 1a). Constituent members of the cell are inclined to the x-y plane at 45◦ which is the minimum angle for self supporting members as observed by Thomas [2]. The structural stiﬀness of the BCC cell along the z axis improves by adding four vertical struts to the cell. The struts connect upper corners of the cell to the lower and transforms the topology into a BCCz cell as seen in (Fig. 1b). The BCCz cell is self-supporting only if the angle between the build platform and all constitutent members is greater than 45◦. The gy- roid cell shown in Fig. 1c represents a minimal surface proposed by Schoen [10].

The gy- roid cell shown in Fig. 1c represents a minimal surface proposed by Schoen [10]. For a three dimensional Cartesian coordinate space x, y, z, the gyroid surface conforms to the equation:sin(2π Lx)cos(2π Ly) + sin(2π Ly)cos(2π Lz) + sin(2π Lz)cos(2π Lx) − t = 0where L is the size of the cube and t is a constant parameter. t assumes a value between 0 and 1.413. A variant of the gyroid, shown in Fig. 1d is the matrix form of the double gyroid (D-gyroid) [11]. The equation for this cell is derived from Eqn. 1 by squaring it’s three terms. Eqn. 1 becomes:2π L F CC includes members that connect the corners of the cube to the center of faces perpen- dicular to the x-y plane. P F CC is a variant of F CC with four vertical struts to improve stiﬀness along the z axis.

P F CC is a variant of F CC with four vertical struts to improve stiﬀness along the z axis. F2BCC is a boolean combination of BCC and F2BCC.2π L2π L2π L2π L2π Lx))2 − t2 = 0(sin(x)cos()y + sin(y)cos()z + sin(z)cos(1239(1)(2)(a)(b)(c)(d)yzx(e)(f)(g)Figure 1: Seven self-supporting unit cells: (a) BCC (b) BCCz (c) Gyroid (d) matrix phase of D-gyroid (e) F CC (f) P F CC (g) F2BCC.3 Finite Element Convergence StudyPreliminary studies were performed to determine cellular parameters require to set the relative density of STL ﬁles at 0.23. The relative densities of BCC, BCCz, F CC, P F CC and F2BCC were controlled by the radius, R, of the cylindrical members. Mem- bers within a cell were constrained to share the same R which was then varied iteratively until the target density was achieved. The resultant values for R are shown in Table 1. R is clearly dependent on the number of members, M , and the length of these members.

The resultant values for R are shown in Table 1. R is clearly dependent on the number of members, M , and the length of these members. For the gyroid lattices, t was varied in Eqn 1 and 2 until the relative density constraint was achieved. A lower value of t was determined for D-gyroid than that for the gyroid as shown in Table 1. For the ﬁrst set of FEA simulations, STL models of the unit cells were voxelized and converted to hexahedral meshes. Voxelization is the conversion of a model represented by surfaces or patches into a bitwise three dimensional array. The array contains the value one in regions enclosed by the STL and zero in unoccupied re- gions. Methods for voxelizing STL models and other triangular facets can be found in [12, 13, 14]. Each voxel enclosed an integral portion of the cells domain with a shape that largely resembles that of a hexahedral element.

Each voxel enclosed an integral portion of the cells domain with a shape that largely resembles that of a hexahedral element. Therefore, regions of the voxel array holding a value of one were mapped directly into an hexahedral mesh as illustrated in Fig. 2 for the D-gyroid cell. The resolution of the voxels was used to control the mesh size during convergence studies. By increasing the resolution of voxels, more hexahedral elements are introduced in the mesh until convergence was reached. Convergence was achieved once the change in maximum displacement along the loading axis was less than 1% for 100% increase in the number of elements.

Convergence was achieved once the change in maximum displacement along the loading axis was less than 1% for 100% increase in the number of elements. The meshing strategy preserved the quality of elements and avoided numerical errors associated with distorted elements.1240Table 1: Parametric values R and t for the unit cells used to set relative density to 0.23.Unit Cell BCC BCCz Gyroid D-gyroid F CC P F CC F2BCCR (mm) t M1.15 - 81.08 - 120.91 -0.36 -1.29 - 161.20 - 200.86 - 24A voxel subsequently converted to a hexahedral elementSTLVoxel ArrayFigure 2: Conversion of D-gyroid STL to voxel array.zyxPlateyzx=0xyz=0xyz=101mm10N/mm2Figure 3: Pressure loading on a D-gyroid cell.1241A set of hexahedral elements constituting a 1 mm thick plate was placed on the cells along the z-axis, on Plane xyz=10. The plate experienced a tensile pressure of 10 N/mm2 on it upper face as illustrated in Fig. 3.

The plate experienced a tensile pressure of 10 N/mm2 on it upper face as illustrated in Fig. 3. Nodes on Plane xyz=0 were ﬁxed for both translational and rotational degrees of freedom. The Young’s modulus of elements associated to the unit cell was set to 100 GPa and Poission ratio was 0.342 which typical for SLM compatible metals. Plate elements assumed material properties that was many times stiﬀer than the cell elements. A static linear structural analysis was performed and the mesh was progressively reﬁned until convergence was achieved. The plate was later rotated so that the pressure on it’s face aligned with the x-axis. Nodes on the Plane yzx=0 were ﬁxed for all degrees of freedom and the lower face of the plate was located at Plane yzx=10 which is the farthest face of the cell along the x-axis.

Nodes on the Plane yzx=0 were ﬁxed for all degrees of freedom and the lower face of the plate was located at Plane yzx=10 which is the farthest face of the cell along the x-axis. The convergence simulations was then repeated for this boundary condition.A second set of simulations was performed to investigate if observed diﬀerences across the cells translates to their lattices. Lattice structures were generated for the cells by tessellating their voxel arrays as proposed in [15]. The length, width and height of each lattice was set to 40 mm while the unit cell size remained at 10 mm resulting in a 4 by 4 by 4 lattice. This is illustrated for the D-gyroid in Fig. 4. To improve computational eﬃciency, a quarter of the lattice was modelled and symmetrical constraints were imposed along the planes of symmetry.

This is illustrated for the D-gyroid in Fig. 4. To improve computational eﬃciency, a quarter of the lattice was modelled and symmetrical constraints were imposed along the planes of symmetry. This is an established way of reducing large FEA model which gives the same results as the full model.yzx=20xyz=0xzy=20Double gyroid lattice40mmReduced modelyzxPlate10N/mm2Figure 4: Constructing reduced FEA models for the D-gyroid lattice.1242This was achieved by diving the full model along centric Plane yzx=20 and xzy=20 and elements in one of the four quadrants were retained. Nodes residing on yzx=20 were ﬁxed along x-axis while those on the xzy=20 were ﬁxed along the y axis. Nodes on xyz=0 were ﬁxed for all degrees of freedom and the upper face of the plate experienced a pressure of 10 N/mm2.

Nodes on xyz=0 were ﬁxed for all degrees of freedom and the upper face of the plate experienced a pressure of 10 N/mm2. Simulations were repeated for pressure loads along the x-axis with sectional Plane xyz=20 and xzy=20 used to reduce the FEA model.4 ResultsDisplacement along z and x axes due to pressure loads are plotted against the number of elements in Fig. 5. Fig.

5a shows results for z-axis while those for x-axis are shown in Fig 5b.F2BCCF2BCCBCCBCCzPFCCFCC00.010.020.030.040.050.060.070.080.000.250.500.751.001.251.50z-Displacement (mm) Elements (million)GyroidD-Gyroid(a)GyroidBCCFCC0.010.020.030.040.050.060.000.250.500.751.001.251.50x-Displacement (mm) Elements (million)F2BCCPFCCBCCzF2BCCD-Gyroid(b)Figure 5: Finite element convergence plot for the unit cells (a) z-displacements (b) x- displacements.1243The following is deduced from the scattered plots:The gyroid cell reached the greatest displacement for both the z and x axes sug- gesting it the weakest of the cells,BCC has the second largest displacement along the z-axes but lower than three other cells along the x-axis,The D-gyroid cell achieved the lowest displacement along the x and competively lower displacement along the z-axisDisplacement for F CC and P F CC cells along both axis are lower than most of the other cells,Convergence plot for the Lattices are shown in Fig.

Comparing the scattered plots for the lattices with those for the cells, it is observed that:D-gyroid lattice achieved the lowest displacement in the x-axis,Displacement for F CC and P F CC cells along both axis are lower than most of the other cells,The D-gyroid lattice has similar stiﬀness to that of BCCz in the z-axis but ap- proximately two times stiﬀer than BCCz in the x-axis,The gyroid lattice is better than the BCC lattice in x and z axes,The ﬁrst three observation are consistent with those of the cells, however, the fourth contradicts deduction from the comparision of the BCC and gyroid cells. This suggests that the properties from both cells is insuﬃcient to determine the performance of the BCC and gyroid lattices.

This suggests that the properties from both cells is insuﬃcient to determine the performance of the BCC and gyroid lattices. Also, the presence of the vertical members in the BCCz only improves the stiﬀness of the lattice along the z and has no eﬀect in the x since BCCz is weaker than BCC along the x-axis.5 DiscussionThe stiﬀness, K of the unit cells and lattices was determined as the ratio of total forceon the plate to the maximum displacement along an axis. Mathematically,K =P A dwhere P is the pressure on the plate and assumes a value of 10 N/mm2 and A is the surface area of the plate. d is the converged maximum displacement along either of the two axis. For the cells, A is 100 mm2 while for the lattice it is estimated to be 1600 mm2. K along z and x axes for cells and lattices are plotted as bar charts and shown in Fig.

For the cells, A is 100 mm2 while for the lattice it is estimated to be 1600 mm2. K along z and x axes for cells and lattices are plotted as bar charts and shown in Fig. 7.020406080100120BCCBCCzD-gyroidFCCPFCCGyroidF2BCCK (N/mm X 10-3) Unit cellzxF2BCCxF2BCCz051015202530BCCBCCzDgyroidFCCPFCCGyroidF2BCCK (N/mm X 10-3) Lattice type(a)(b)Figure 7: Maximum stiﬀness for the cells (a) Lattice (b).1245(3)BCC, D-gyroid and gyroid cells and lattices have approximately similar K along both z and x axes. This could be attributed to their cubic symmetric properties. The F2BCC lattice also show similar behaviour in both z and x axes (Fig. 7b), this is not evident for the F2BCC cell (Fig. 7a). Also, BCCz, F CC, and P F CC showed contrasting stiﬀness in the axes. Observed peak stiﬀness for the lattices is exhibited by P F CC along the z−axis which is a consequence of the minimal displacement in this axis.

Observed peak stiﬀness for the lattices is exhibited by P F CC along the z−axis which is a consequence of the minimal displacement in this axis. However, P F CC stiﬀness along the x-axis is approximately 30% lower than that for D-gyroid. The maximum von Mises stresses for cells and lattices are shown in Fig. 8.F2BCC020040060080010001200140016001800BCCBCCzD-gyroidFCCPFCCGyroidF2BCCMaximum Stress (MPa) Unit cellzx0200400600800100012001400160018002000BCCBCCzDgyroidFCCPFCCGyroidF2BCCMaximum Stress (MPa) LatticezxF2BCC(a)(b)Figure 8: Maximum stresses for cells (a) lattices (b).It’s variant, the The D-gyroid lattice and cell exhibit the lowest stress in both axes. gyroid, performs the worst achieving stresses many times greater than that of the D- gyroid. Again, both D-gyroid and BCC show similar stress levels in both axes, however, maximum stresses in the D-gyroid are a fraction of those in BCC.

Again, both D-gyroid and BCC show similar stress levels in both axes, however, maximum stresses in the D-gyroid are a fraction of those in BCC. Stress levels in P F CC and F CC along the z-axis is better that in the x-axis. P F CC and F CC was seen to have greater stiﬀness along the z-axis than those of D-gyroid, however, stresses in their cell and lattice are much larger as seen Fig. 8a and 8b.For structural applications involving pressure loads along a single axis, F CC or P F CC, is better than the other ﬁve types. It is common for components to experience multiple pressure loads in diﬀerent orthogonal axes, it is therefore preferrable to utilize a lattice that shows superior performance in more than one axis. For this instance, the D-gyroid lattice is better than the other lattices since it acheived competitive stiﬀness in both z and x axes with lower stresses.

For this instance, the D-gyroid lattice is better than the other lattices since it acheived competitive stiﬀness in both z and x axes with lower stresses. The behaviour of the cells and lattices have not been determine for pressure loads along the y-axis since this can be deduced from observations in z and x axes. BCC, BCCz, F CC, P F CC and F2BCC are symmetrical about centric zy planes, implying that their behaviour along the y-axis is largely similar to those in the x-axis if the similar pressure loads were applied in both axes. Both gyroid and D-gyroid are not symmetric about the centric zy planes, however, they do exhibit cubic symmetric properties as described by Scherer [11]. It was seen that their stiﬀness is largely similar in both z and x axes. Therefore, properties of both lattices along y-axis would be similar to those observed in z or x.

It was seen that their stiﬀness is largely similar in both z and x axes. Therefore, properties of both lattices along y-axis would be similar to those observed in z or x. Superior qualities of the D-gyroid lattice automatically translates to the y-axis.1246Thermostructural applications of lattices are common. These have not been consid- ered in this work, however, the relative thermal dissipative and insulative properties of the cells is in principle deducible from their total surface areas. The surface areas of the cells are plotted in Fig. 9.F2BCC3.003.504.004.505.005.506.006.507.007.50BCCBCCzD-gyroidFCCPFCCGyroidF2BCCSurface Area (X 10-2 mm2) Unit CellsFigure 9: A comparison of surface area of the unit cells with similar density.It can be seen that the D-gyroid lattice has the greatest surface area, suggesting it has the greatest capacity to dissipate heat when a cooling agent is ﬂushed across it’s surface.

F2BCC also show some capacity to dissipate heat in since it has a greater surface area than either BCC, BBCz, F CC, P F CC and gyroid; therefore, there is greater interaction between the agent and F2BCC surfaces. For applications where thermal insulation is paramount, the BCC and the gyroid lattice are better since they have the least capacity to transport heat owing to their relative low surface areas. However, surface defects would propagate faster on D-gyroid and F2BCC since they include thinner members in a broader space.6 Summary and ConclusionsThe capabilities of SLM is evident for metallic components since a wider range of topolo- gies can be realized as opposed to traditional techniques. However, most would require support structures to prevent warping and curling of the part during the build, while other would self-support. Diﬃculties associated with removing these supports makes self-supporting solutions more suited for SLM.

Diﬃculties associated with removing these supports makes self-supporting solutions more suited for SLM. Supports are particularly diﬃcult to re- move from lattices, therefore it is preferrable to design latticed components that reduce the need for supports. Interestingly, a number these lattices exist, some of which have pre- viously been characterized with SLM. In this work, the structural stiﬀness and stresses in BCC, BCCz, F CC, P F CC, D-gyroid, F2BCC and gyroid cell and lattices were compared. To ensure a fair comparison is made across the lattices and cells, the relative density for cells was ﬁxed at 0.23. The F CC, P F CC and D-gyroid lattices were found to exhibit superior perfomance to those of the other lattices. Also, the D-gyroid cell1247has the greatest surface area implying that it has a better potential to dissipate heat if it’s interact with a cooling agent.

Also, the D-gyroid cell1247has the greatest surface area implying that it has a better potential to dissipate heat if it’s interact with a cooling agent. BCC and gyroid lattice are poorest at dissipating heat owing to their relatively low surface area and can potentially be used as a thermal insulator.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceA Comparative Investigation of Sintering Methods for Polymer 3D Printing Using Selective Separation Shaping (SSS)H. Nouri*, B. Khoshnevis*Department of Industrial and Systems and Engineering, University of Southern California, Los Angeles California, 90089AbstractSelective Separating Shaping (SSS) is a novel additive manufacturing process which is capable of processing polymeric, metallic, ceramic and cementitious materials. In earlier experiments, the capabilities of SSS in fabrication of metallic, ceramic, cement-based and polymeric parts have been demonstrated. The focus of this research has been on exploration of sintering methods in SSS for successful fabrication of polymeric parts. The SSS machine has been used to build specimens made of polyamide (PA6) material.

The SSS machine has been used to build specimens made of polyamide (PA6) material. Bonds between layers under two different thermal sintering methods are investigated to achieve better control over shrinkage and maintain effective binding between layers. ImageJ platform and binary surface plots have been used for image processing and evaluating final porosity under each heating mechanism. Further investigations are carried out on properties of the base materials and the choice of sintering mechanism to further improve resolution of final parts.IntroductionAlthough AM has developed rapidly and is receiving increasing attention, it is still in early development stage and still significant amount of improvements are expected for AM machines to become mainstream manufacturing equipment. Improvements are expected in the following direction: i. Increasing speed of printing by maintaining quality of printed parts, ii.

Improvements are expected in the following direction: i. Increasing speed of printing by maintaining quality of printed parts, ii. Online monitoring of process parameters and in-situ sensing techniques to identify deviations and correct them during the process, iii. Apply new materials with desired chemical properties and microstructures. iv. Reducing waste material and further research on reusing recycled materials. v. Fabrication of multi-functional parts and control of material properties and their performance. vi. New AM technologies which can process a wider range of materials in a single platform without a need for retooling. To further expand application of AM to functional engineering parts rather than prototypes, several challenges need to be addressed. There is high demand for a technology that reduces time-to-market and can produce lower cost parts while maintaining high quality.

There is high demand for a technology that reduces time-to-market and can produce lower cost parts while maintaining high quality. Additionally, a machine with the flexibility in applying multiple classes of material can be more affordable for a firm than multiple machines that each process a certain class of materials. Selective Separation Shaping (SSS) was developed with the goal of fabricating low cost, high resolution parts in a single multi-material platform. This process is not only limited to small scale fabrications, but it also has potential for fabricating on demand large scale structures. For polymeric part fabrication, due to the operation principle in SSS (printing on the boundaries instead of the part body), this process can compete with other processes such as FDM and SLS in printing speed. Multiple nozzles can print at once on the platform without interfering with one another to build multiple objects simultaneously.

Multiple nozzles can print at once on the platform without interfering with one another to build multiple objects simultaneously. Compared to other powder-based polymer processes, the end-user product cost will be low as this process does not require laser scanning. The separating material can be purchased at low cost. Additionally, a variety of polymers can be applied to the machine and be tested in powder form. Table 1 summarizes the AM 3D printing fabrication methods for polymer fabrication.1474Table 1.

Additionally, a variety of polymers can be applied to the machine and be tested in powder form. Table 1 summarizes the AM 3D printing fabrication methods for polymer fabrication.1474Table 1. AM methods for fabrication of polymer part geometry creation methodsAM technology3D printing methodSLAUV light beam/LaserSLS FDM 3DP DLP DMLS SHS SLM SISLaser Extrusion Binder Digital projector screen Laser Heat Laser Separator deposition (Liquid)SSSSeparator deposition (Solid)In earlier experiments [1], sample polymer parts with multiple overlapping layers were fabricated by SSSusing Copolyamide and Vestamelt 171.Table 2 shows a summary of optimum process parameters used.Table 2.

Summary of process parametersProperties Thickness of each layer Sintering speed Sintering temperatureDescription 0.09 to 0.25 mm 8.3 mm/sec 65-70 CPrinting speed Average wait time between layers 12 Sec Build tank temperature Platform temperature Chamber temperature Type of polymers Powder size range S-powder size range32 sec/layer48 to 60 C 54 C 30 C Nylon 6,6 90 to 250 micrometers 25 to 45 micrometersPrevious SSS research has offered insight into many unknown parameters in the SSS process [2-5].Zhang et al focused on high temperature sintering material and successfully fabricated metal and ceramic parts. However, bulk sintering approach used in SSS for metal and ceramic has limitations and is not applied to polymers. Therefore, layer by layer sintering approaches need to be carefully studied and implemented.Weak bonding between layersOne of the key factors in successful fabrication of polymeric parts with SSS is effective sintering of layers.

Increasing layer thickness can result in inadequate transfer of heat energy for diffusing successive layers. Figure 1 shows poor bonding between layers for 200 micrometers layer thickness. Depending on the thickness of the layer, sintering adjustments should be made to achieve effective bonding between layers.1475Figure 1. Partially sintered sample with weak bonding between layersCumulative heating and warpageDepending on the properties of base material and sintering and cooling temperatures and rates, warpage of the layers can occur as shown in Figure 2. Also, as the number of layers increases this phenomenon becomes more significant. Warpage can occur by penetration of heat into previously sintered layers resulting in over- sintering. Moreover, depending on the melting temperature (Tm) and crystallization temperature (Tc) this warpage can be eliminated in polyamide samples with smaller window between Tm and Tc [6-8].Figure 2.

Moreover, depending on the melting temperature (Tm) and crystallization temperature (Tc) this warpage can be eliminated in polyamide samples with smaller window between Tm and Tc [6-8].Figure 2. Warpage caused by cumulative heatingNon-uniform heating and cooling can be other reasons for layer warpage. Therefore, attempt is made tokeep the temperature inside the chamber at a constant level and to pre-heat the platform before printing starts.Shift between layersIn successive deposition of layers, it is very important to have control over uniformity in the deposition of B-powder over sintered layer to avoid any displacement of previous layer. Shift between layers is one of the main causes for bad quality in the final part.

Shift between layers is one of the main causes for bad quality in the final part. Shifts can be caused by non-uniform sintering of layers, collision of roller with a sintered layer due to inadequate distance from the layer, high friction between roller and surface of powder in build platform, previous layer’s warped edge which is easily hit by the roller, software failure, and poor bonding between successive sintered layers. Moreover, B-powder’s flowability, particle size and shape as well as roller (or blade) properties such as shape and material, are important factors in successful spreading of powder on previously sintered layer and need to be considered. Finally, Excessive amount of powder fed to the build tank was also observed to be a contributing factor to layer shifting.ShrinkageMaterial shrinkage is one of the main sources of part inaccuracy in AM, primarily caused by thermal phase change of the material during layer formation [9-11].

Proper adjustment of process parameters can severely affect the output quality of the products. In terms of process planning, many research efforts have been dedicated to look for the relationship between the output part quality and the input building process parameters [12-15]. Parametric process optimization is used to improve functional requirement of an additive manufacturing system 1476such as accuracy by finding the optimal settings of factors such as light intensity of the machine, laser beam width, light exposure duration, scan speed, layer thickness, and so forth. It is crucial to find the appropriate process parameter settings which lead to the best trade-off among part accuracy and manufacturing build time. Several studies have been dedicated to improving part quality by modifying process variables. Sager et al. [16] studied a stereolithography process and investigated the relationship between laser scan speed and surface finish of the final part.

Sager et al. [16] studied a stereolithography process and investigated the relationship between laser scan speed and surface finish of the final part. He concluded that appropriate laser scan speed leads to a better surface finish of SLA geometries. Raghunath and Pandey [17] investigated the relationship between process parameters and material shrinkage. They examined laser power, beam speed, hatch spacing, part bed temperature and scan length and adopted Taguchi method for analysis of shrinkage. It was concluded that laser power and scan length are found to be most significant process variables influencing shrinkage in the X direction. Senthilkumaran et al. [18] also focused on shrinkage behavior by analyzing the relationship between shrinkage calibration specimen and various building conditions in SLS parts.

Senthilkumaran et al. [18] also focused on shrinkage behavior by analyzing the relationship between shrinkage calibration specimen and various building conditions in SLS parts. In order to reach an accurate estimation, layer thickness, laser power, hatch spacing, scanning speed, part bed temperature and scanning mode were studied as the process parameters set in the experiment and beam offset and scanning position were observed to have the major influence on the shrinkage of the part. The focus of this section is to address shrinkage in SSS using process planning and machine parameter setup adjustment. The following section presents the study performed on the strength of the bond between layers using forced convection and the corresponding results.Sintering and coalescencePowder sintering refers to a thermal process used to convert a powder volume into a solid object. Polymer sintering is the formation of homogenous mass from the diffusion of smaller particles.

Polymer sintering is the formation of homogenous mass from the diffusion of smaller particles. Generally, this process is carried out between the glass transition temperature (Tg) and melting temperature (Tm) of a specific polymer. When polymer particles are heated above Tg, they form a neck at which polymer chains can interact and diffusion occurs. When the entanglements are fully established, this mass will be at equilibrium [19]. Figure 3 shows growth of the contact neck between powder particles.Figure 3. SEM pictures of powder before(A) and after sintering(B)Heat can be transferred between objects by means of radiation, convection, conduction, and microwave[20]. Extensive research work has been done in developing a thermal analysis method for powder metallurgy processes such as Laser sintering [21-23]. Uniform temperature distribution in powder sintering processes can severely affect the final quality of the part.

Uniform temperature distribution in powder sintering processes can severely affect the final quality of the part. Understanding process mechanisms can significantly contribute to the future development of SSS parts. In the following section, two methods of powder sintering that were examined for SSS of polymeric material are discussed. In the following section two methods of powder sintering that were examined for SSS of polymeric material are discussed.Experimental studyIt is hypothesized that non-deformed inter-layer adhesion can be achieved by convection-based heatingat a speed which is at least the same as speed of radiation-based sintering which can produce the same result. Investigation are carried out on base materials properties and the choice of sintering mechanism to further improve mechanical strength and achieve acceptable surface quality. Microanalysis of layers have been 1477performed to obtain a better understanding of the process.

Microanalysis of layers have been 1477performed to obtain a better understanding of the process. The impact of infrared and convection sintering on samples have been studied. To have a better comparison of the two systems, the amount of heat applied to the powder by each method is determined by measuring the temperature at the point where the powder is sintered. Vestosint material has been used as the base powder for the first set of experiments. Sintering speed is kept constant during all experiments. To study the impact of heat flow on sintered layers, a set of experiments were conveyed with variant distance of convection heater to the powder. Figure 4 represents a binary surface plot of layers sintered under convection heating from distances of 11mm and 4 mm from surface of powder. These distances are selected such that 11 mm is the maximum distance from which sintering is possible and 4 mm is the shortest distance possible to the surface of the powder.

These distances are selected such that 11 mm is the maximum distance from which sintering is possible and 4 mm is the shortest distance possible to the surface of the powder. These plots are created based on Gaussian light distribution of microscopic images. The heat intensity has been constant in both cases. Although in case b, there is a greater amount of heat flowing inside the powder relative to higher distance exposure, it is observed that decreasing distance of the heater from powder has increased porosity on the layer due to high velocity of the flow. Is it observed that in both cases heat penetration depth is relatively high which indicates that in cases where high volume sintering is desirable this approach is suitable (Figure 5).Figure 4. Surface plot of convection sintering at 11mm distance (left) versus 4 mm (right) from surface layerFigure 5.

Surface plot of convection sintering at 11mm distance (left) versus 4 mm (right) from surface layerFigure 5. SEM image of Vestosint at 150 magnitudes sintered at distance 11mm(left) and 4mm (right)Increasing the heat intensity under the same condition of distance from surface of powder and flow velocity, results in stronger bonds between particles while maintaining average pore size diameters the same, as opposed to the previous condition in which average pore size was increased. In Figure 6 the surface of Vestosint under low (left) and high (right) heat intensity is shown.Figure 6. Surface of Vestosint under in low (left) and high level heat intensity1478In the case of radiation, it is observed that increasing heat intensity from low to high level results in complete coalescence of particles on surface (less porosity on surface); however, does not guarantee bottom particle diffusion (Figure 7).Figure7.

Surface of Vestosint under IR heating in low and high-level heat intensity (4 mm distance)In the case of convection sintering, even at high level heat intensity, the energy is absorbed by all particles causing sintering from the bottom up through the whole body. While radiation heat is absorbed in the first layer, any remaining energy is transferred by conduction from top to bottom gradually (Figure 8).Figure 8. Surface plot of Vestosint under Radiation heating in low and high level heat intensity (4 mm distance)Bright sections on these images are indicative of light transfer through the layers and therefore, sintering level.

Similarly, in both cases, bottom layers have a higher level of density as moving up to upper layers shows a decreasing sintering ratio.Selective sintering by forced convective sintering methodWith an understanding of sintering behavior of base material, it was necessary to determine whether or not an effective layer bonding could be achieved through other heat transfer mechanisms. Based on preliminary experiments by linear convective heater discussed in previous section, it was hypothesized that non-deformed inter-layer adhesion can be achieved by convection-based heating at a speed which is at least the same as the speed of radiation-based sintering which can produce the same result. Interlayer adhesion of the top two layers is achieved primarily by conduction of heat from top to lower layer. Conduction takes times, but faster conduction may be achieved with higher radiation intensity but at the expense of layer deformation.

Conduction takes times, but faster conduction may be achieved with higher radiation intensity but at the expense of layer deformation. More suitable layer shape may be achieved at lower heat delivered over extended time to allow heat conduction from higher to lower layers. However, slow sintering prolongs part fabrication time. Convection-based heating can penetrate heat into top two layers through the pores of the top layer and sinter the two layers simultaneously without the delay that conduction heating would require. The first step was to build parts using this heat transfer approach.1479Figure 9. Schematic of forced convective heaterA pointwise heater with outer diameter of 6mm and inner diameter of 4 mm was used for experiments. Main components of heater include ceramic tube, kanthal coil, as well as air pump which blows air through the coil along the ceramic tube.

Main components of heater include ceramic tube, kanthal coil, as well as air pump which blows air through the coil along the ceramic tube. A heater cover plate module is installed and is used to block and unblock heat exposure when desired (Figure 9). It was believed that switching heater between on and off state modes on each layer is associated with delays in reaching desired temperature. Thus, week degree of bonding would be achieved at start points as well as undesired sintered areas following end points.Figure 10. Demonstration of sintering by forced convective heaterUpon completion of deposition on a layer, the base plate moves down by an increment equal to half of the layer thickness. Heater is moved over the surface as it can be seen in Figure 10. Sintering starts as the heater moves along the boundary and inner body areas (Figure 11).1480Figure 11.

Heater is moved over the surface as it can be seen in Figure 10. Sintering starts as the heater moves along the boundary and inner body areas (Figure 11).1480Figure 11. Sintering profile around the contour of the partWhen sintering is completed, the cover plate module is closed at the endpoint and the base plate movesdown by an increment equal to half of layer thickness and the next layer is deposited.DOE for determining significant factors in forced convective sinteringDesign of experiments allows better understanding of relationships between the inputs and the response. A design of experiment technique is used to test the experimental factors at different levels. The motivation to use this approach in this investigation is that it is highly efficient for analyzing process parameters and their interactions effect in convective sintering. This experiment is especially important as it is indicative of the bond formation between successive layers.

This experiment is especially important as it is indicative of the bond formation between successive layers. The design matrix can be seen in Table 4.Table 3. FactorsA: Distance from powderB: Heat intensityC: Air flowD: Speed of movementThis experiment was designed for the following purpose: 1. The interaction between distance of the heater from the surface, air pressure of the pump, heat intensity and finally speed of sintering are studied. Optimal setting can be found accordingly for more effective sintering. 2. Effective layer bonding can be examined through evaluation of layer thickness and strength of specimens, 3. Heat transfer mechanisms are compared, and better understanding is obtained. Levels are decided such that, in case of high distance from the surface, it is assumed the main source of sintering is forced convection while in case of low distance and low air pressure, radiation is expected to be the main source.

It is important to note that in all cases, there is heat transfer through conduction. However, it is assumed it has been kept equally the same among all runs.1481Table 4. Design matrix (1 at high level and -1 at low level)TrialDistance from powderHeat intensityPump intensitySpeed of movement1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 161 -1 -1 -1 -1 -1 -1 -1 1 1 1 1 1 1 1 11 -1 -1 -1 1 1 1 1 -1 -1 -1 -1 1 1 1 11 -1 1 1 -1 -1 1 1 -1 -1 1 1 -1 -1 1 11 1 -1 1 -1 1 -1 1 -1 1 -1 1 -1 1 -1 1For each of the trials shown in the matrix at Table 4, a path with a length of 40 mm and diameter of the heater was sintered. The width of the paths varies depending on distance of the heater from the surface as well as air pressure. Each run was repeated 3 times and to reduce bias experiments were done in a randomized order. The average width, thickness, and tensile strength of each specimen was recorder and are reflected in Table 5.Figure 12.

Comparing values, highest layer thickness and strength was achieved in runs 7 and 3 respectively which suggest high air pressure, low sintering speed. Half-normal plot shown in Figure 13 represent significant factors and interaction in determining layer thickness. As it can be understood, the interaction between air pressure and speed plays a more important role in determining layer thickness rather than interaction between heat intensity and speed.Figure 13. half normal plot for layer thickness1483It is observed that geater layer thickness shows better heat delivery to bottom layers. Higher penetration of heat through pores is achieved by convective heating. Effective heat transfer results in homogeneous densification and affects microstructure of sintered samples.DiscussionSuitable layer shape may be achieved at lower heat delivered over extended time to allow heat conduction from higher to lower layers. However, slow sintering prolongs part fabrication time.

However, slow sintering prolongs part fabrication time. Convection-based heating can penetrate heat into top two layers through the pores of the top layer and sinter the two layers simultaneously without the delay that conduction heating would require. In our future research, sample parts will be built and tensile tests will be performed to examine mechanical properties of samples.AcknowledgementSSS has been conceived and developed in the course of a Phase-II NASA Innovative Advanced Concepts (NIAC) project in relation to the part of the project that concerned landing pad construction.1484References[1] Nouri, H., and B. Khoshnevis. "Selective Separation Shaping of Polymeric Parts." In Solid Freeform Fabrication Symposium. 2017.[2] Khoshnevis, Behrokh, and Jing Zhang. "Selective Separation Sintering (SSS)-An Additive Manufacturing Approach for Fabrication of Ceramic and Metallic Parts with Application in Planetary Construction." AIAA SPACE 2015 Conference and Exposition.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA Comparative Study Between 3-Axis and 5-Axis Additively Manufactured Samples and their Ability to Resist Compressive LoadingN.Kaill*,R.I.Campbell*, P.Pradel*, G.A.Bingham+Loughborough Design School, Loughborough University, United Kingdom +De Montfort University, United KingdomAbstractOne of the main limitations of parts made with Material Extrusion (ME) is their anisotropic mechanical behaviour. This behaviour limits the functionality of these components in multi-directional loading conditions. A critical factor for this mechanical behaviour is the poor bonding between layers. 5-axis ME has the capability to orientate the printed layers in order to limit the effect of poor inter-laminar bonding.

5-axis ME has the capability to orientate the printed layers in order to limit the effect of poor inter-laminar bonding. Previous studies have investigated 5- axis ME, but not fully explored 5-axis capabilities of this manufacturing technique. To address this gap, this paper compares the mechanical behaviour of 3-axis and 5-axis ME samples when subjected to compressive loading. The results demonstrate how depositing material in “3D layers” can improve the consistency of a sample’s mechanical behaviour. This study indicates that 5-axis ME can enable more isotropic behaviour in printed samples.IntroductionMaterial extrusion (ME) is the most prevalent category within additive manufacturing (AM) (1). ME is the process by which a material, in a liquid state, is metered through a nozzle or orifice (2). The extruded bead of material meets the build plate or previous layer and adheres to it.

ME is the process by which a material, in a liquid state, is metered through a nozzle or orifice (2). The extruded bead of material meets the build plate or previous layer and adheres to it. Typically, the extruder head is moved in the x-y axes via a gantry system that allows the creation of the layer detail. Once the layer is completed either the printed head is raised, or the build plate lowered by the layer thickness. This accumulation of multiple layers builds three- dimensional geometries. From a mechanical perspective, ME components present an anisotropic mechanical behaviour, because the inter layer bonding is much weaker than the intra layer one (3–8).According to the 2017 Wohlers report, functional parts are the largest single application for AM produced components, with 33.8% of AM printed parts created for this purpose (9). For ME parts to be utilised for functional end-use applications, they must withstand complex multi-directional loadings.

For ME parts to be utilised for functional end-use applications, they must withstand complex multi-directional loadings. To address this limitation, the authors speculated that 5-axis ME could be utilised to improve parts’ mechanical behaviour by aligning the filament in the direction of the loads. In a previous paper, it was established that multi-axis ME can indeed improve the mechanical behaviour of ME parts (10). However, that study was limited to 4.5 axes due to one of the rotational axes being locked perpendicularly. This limitation was suitable for the geometrical complexity of the examined parts, but it did not address the difference in mechanical behaviour between fully 5-axis and 3-axis samples.

This limitation was suitable for the geometrical complexity of the examined parts, but it did not address the difference in mechanical behaviour between fully 5-axis and 3-axis samples. Similarly, other studies that have looked into curved layer deposition have utilised a DELTA parallel robot (5,11) resulting in parts that were limited by the fixed position of the nozzle.This paper will expand on previous work by exploiting 5-axis capabilities through a highly curved geometry represented by a simple thin-walled dome. Due to its topology, this1818geometry is appropriate to explore how to optimise the mechanical performance of ME parts by exploiting 5 axes capabilities to selectively orient the deposition of filament.The geometry that was the selected for this study is a dome with a interal radius of 15mm and a wall thickness of 3mm.

This geometry was selected as it is a simple shape that would allow for 5-axis ME to be explored with the revlevent directionality associated with this AM method. The thickness was selected based on the number of shell that could be used via a 5-axis method, with seven shells being decided upon. With the internal radius set at 15mm to keep samples in line samples produced in previous studies using similar methods (10). Traditional compression sample may be able to produce cleaner results, but the geometry may restrict the investigation into full 5-axis ME. This restriction could impede the comparrision between three and five axis ME.MethodA finite element analysis (FEA) simulation was conducted to determine the direction of the stress vectors within the geometry of the dome when loaded in compression in the direction of its vertical axis (Figure 1). The simulation was carried out within Solidworks CAD software (12).

The simulation was carried out within Solidworks CAD software (12). A section of the dome’s geometry was used in the simulation, with the boundary surfaces being constrained to allow sliding. A force was then applied to the top of the section, to induce the resultant stress vectors. The simulation indicates that the vertical direction is the dominant direction of the stress vectors. The horizontal stress component was indicated to be the next dominant component. In Figure 1, the colour of the vectors denotes the magnitude of the stress, in Solidworks red is used to indicate high values, while green indicates a lower value. When a compressive force of 1000N is applied, the vertical stress vectors have a maximum value of 12.34MPa with the horizontal having a maximum value of 3.84MPA.This information was then used to create printing strategies for the 5-axis samples, as the direction of these stress vectors indicates the direction that material should be deposited in.

As there is no “slicing” software available for converting CAD models to 5-axis tool paths, the tool paths for the experiment must be generated by the researchers. This process started with the print pattern created on a 2D plane, this pattern was then wrapped onto the surface that it will be printed on (i.e. the dome) and offset by half the thickness of the deposited material. Examples of the tool path for the initial layers of the samples are shown in Table 1. These curves are then scaled through parametric modelling techniques in Rhino Grasshopper (13), with each layer being joined to create a single continuous three-dimensional curve. This curve is then converted into a G-code for the 5-axis ME machine using proprietary software provided by the machine vendor (14).1820Table 1: Sample Print StrategiesPrint Strategy 3 Axis5 AxisHorizontalVerticalConcentricCombinedThe next stage was to create the printing jig as the base for the first layer to be deposited onto.

Typical 3-axis ME processes require the use of support structures for printing 3D geometries (15). The printing jig acted as a support for the 5-axis machine and was reused across multiple prints. The 3-axis samples were printed utilising a MakerBot Replicator 2 Machine, whereas the 5-axis samples were printed utilising a 5axismaker machine. With the same material and deposition temperature being used on both machines to maintain consistency.The initial set up period for the 5-axis machine required more time than a conventional 3-axis ME machine. This was due to the need to establish the central axis for each print, as it is entirely user-defined. To reduce the effect of this on the ability to print samples, the jig was1821designed to be modular the purpose of this design choice was to allow the section of jig that the samples were printed on to be removed.

This would ensure that the retained section of the print jig would be left in place maintaining the relative positioning in the machine. The process of depositing material in this fashion is shown in .Figure 3: 5-Axis Material Extrusion: (left) Combined Print, (right) Horizontal PrintAfter printing each sample, its mass and dimensions were measured prior to the compression testing. The dimensions of the samples were measured through the use of a digital calliper, while the weight was measured through the use of digital scales. The average values are shown in Table 2. The printing times included show purely the printing time and do not include any post-processing the samples required. 5-axis samples required minimal post- processing due to the lack of support structures, whereas the 3-axis sample required significantly more time to remove the support structure. Examples of the 3-axis samples are shown in Figure 4.

Examples of the 3-axis samples are shown in Figure 4. This figure illustrates the amount of support material that had to be removed prior to the compression testing.Figure 4: 3-Axis Material Extrusion with Support Structures: (left) Vertical Print, (right) Horizontal PrintAnother interesting point that is highlighted in Table 22 is the difference in mass between the two printing processes, with all the 3-axis and 5-axis displaying similar mass within their own process. To establish fair comparison data sets, the CAD model used for the 3-axis sample was a solid dome with the average wall thickness of 5-axis samples. As the 3- axis samples display a thicker wall than the 5-axis samples, the residual support material may1822be left attached to the surface of the samples.

As the 3- axis samples display a thicker wall than the 5-axis samples, the residual support material may1822be left attached to the surface of the samples. For each data set, 5 samples were produced in order to establish a reasonable average and to make identifying any outliers easier.Table 2: Dome Study Average MeasurementsSampleMean Diameter (mm)Mean Thickness (mm)Mean Mass (g)Print Time (min)3-Axis Horizontal 3-Axis Concentric 3-Axis Vertical 5-Axis Horizontal 5-Axis Vertical 5-Axis Combined35.60 35.93 35.64 35.78 35.86 35.842.88 2.86 3.00 2.74 2.85 2.906.18 6.15 6.20 5.22 5.38 5.3655 53 40 94 73 84The axial compression testing was carried out on an Instron 3366 material test system with a 5kN load cell and a compression speed of 2mm.s-1. The purpose of this test procedure was to establish the different mechanical behaviours of the test samples in regard to their respective stiffness, ductile behaviour and maximum compressive loading.

The purpose of this test procedure was to establish the different mechanical behaviours of the test samples in regard to their respective stiffness, ductile behaviour and maximum compressive loading. The experimental setup is illustrated in Figure 5.Figure 5: Compression Testing Set-up1823Results and DiscussionThe data from the compression testing of each sample were tabulated into an excel spreadsheet, for all 5 samples in each data set. The mean average for each set of samples is displayed in Figure 6.800070006000)5000N( e c r o F d ei l40003000p p A200010000012345671000Displacment (mm)3 Axis Concentric Average3 Axis Horizontal Average3 Axis Vertical Average5 Axis Combined Average5 Axis Horizontal Average5 Axis Vertical AverageFigure 6: Average Compression Test Results for Dome SamplesThe results of this study contain some unexpected findings.

What stands out in Figure 6 is how the 3-axis Horizontal and Concentric samples were able to withstand significantly more load compared to all the other samples. Two possible explanations could be at play. First, the load was applied in one direction to a continuous symmetrical geometry enabling the stacked layers to be in constant compression. This led to the poor inter-laminar bonding not being a major limiting factor until the sample began to deform radially, at which point, the sample failed catastrophically. What is implicated here is that these samples were in compression until failure, meaning the samples behaved more lick a strut than a thin-walled dome. This leads to the second explanation. All the samples followed similar failure behaviour, that of catastrophic failure. This differs from previous studies done by the Izmir Institute of Technology in Turkey that has analysed thin-walled domes (16).

This differs from previous studies done by the Izmir Institute of Technology in Turkey that has analysed thin-walled domes (16). In that work, a dome deforms by first flattening its top, then the centre section is forced into space beneath the dome. Comparing the failure behaviours of the samples within this study the wall thickness is too high to allow for this plastic deformation. The sample shown within Figure 7 shows how much the 3-axis samples typically deformed prior to catastrophically failing. The specific sample shown is from a 3-axis horizontal sample. The top surface was deformed to the point where the centre should be forced through the centre; however, the thickness prevented this type of plastic deformation due to the amount of material present. The 3-axis vertical samples followed this18248failure behaviour but at a significantly lower applied force, compared to all other samples.

The 3-axis vertical samples followed this18248failure behaviour but at a significantly lower applied force, compared to all other samples. An explanation is that the forces where able to shear the interlaminar bonding, with minimal plastic deformation occurring. This confirms how significantly intralayer and interlayer bonding and thus print orientation affect the mechanical behaviour of 3-axis samples (16–19). As to why the 3-axis Concentric and Horizontal samples displayed different values this can be explained by analysing their infill patterns. The concentric 3-axis samples had a concentric infill pattern while the horizontal pattern had a linear infill pattern. The linear nature of the infill pattern allowed it to resist some of the radial deformation allowing for the sample to withstand further loading.

The linear nature of the infill pattern allowed it to resist some of the radial deformation allowing for the sample to withstand further loading. In comparison, the concentric infill pattern could only rely on inter- road bonding with each layer, to resist the radial deformation.Figure 7: 3-Axis Horizontal Sample Post Compression TestingTo understand why our samples’ failure behaviour differs from other work conducted on similar geometries, a sample 3-axis horizontal sample was printed with a wall thickness of 1.2mm, Figure 8 depicts the failure deformation of these samples. Notably, these samples follow the failure behaviour that was expected for this geometry, with initial plastic deformation occurring when a force of 500N was applied to the sample. The outcome of this clarification experiment indicates that there is a wall-thickness boundary condition for this type of geometry.

The outcome of this clarification experiment indicates that there is a wall-thickness boundary condition for this type of geometry. This boundary condition drastically alters the failure behaviour of the samples and the amount of force they can withstand. This result highlights that if the desired geometries are not likely to allow plastic deformation and bending to occur, 5-axis ME may not be the most suitable manufacturing process. This would be an important factor where isotropic behaviour may not be required.Another aspect of the 3-axis sample that may have impacted the results come from the support structures themselves. The support structures used in sample creation where break away supports made of the same PLA material as the sample. This means that the support material and any remaining support material could impact the compression test results.

This means that the support material and any remaining support material could impact the compression test results. Perticularly in regard to the different support structures used to create the 3-axis samples, and how this extra material can alter the compression behaviour of the samples.1825Furthermore, this indicates that there is a further area of research indicated within this study, i.e. to look at samples with smaller wall thickness. A comparison could be made between how 3 and 5 axis samples differ in their compressive resistance when following this different failure behaviour.Figure 8: Thin-Walled 3-Axis Horizontal SampleThe initial shape chosen for this study was selected based upon the assumption that they would follow the failure behaviour similar to those seen in other studies (20) while being a suitably complex shape that could be still be created via ME manufacturing techniques.

A dome represented the simplest shape that incorporated one axis of rotational symmetry and can be printed with 2 axes of rotation. The method for creating the required tool path that enabled 5- axis deposition has highlighted that it is possible to create complex tool paths related to a known surface. This means that more complex geometries that would highlight the benefits of 5-axis could also be created and tested, in order to further explore the potential benefits of 5-axis ME. Focusing purely on the 5-axis samples, the vertical print strategy was able to withstand more applied force than the horizontal print strategy. This does indicate that orientating material to resist the 1st principle stress vector does improve the samples’ mechanical behaviour. Moreover, the combined print strategy was able to resist more force over a prolonged period.

Moreover, the combined print strategy was able to resist more force over a prolonged period. This implies that alternating print patterns that resist a combination of the dominant principle stresses can create samples that display improved mechanical behaviour. This is a potential area of further research that would focus on how much material needs to be deposited along the different dominant stress vectors dependent on their magnitude. This is best explained in Figure 9, which contains a sample created to demonstrate how the alternating shells where deposited. More complicated shell patterns are possible in order to create samples that allow for bending moments to create plastic deformation in a non-catastrophic manner.

More complicated shell patterns are possible in order to create samples that allow for bending moments to create plastic deformation in a non-catastrophic manner. This would allow for better control over failure behaviours and allow samples to behave in a more isotropic manner.1826Figure 9: 5-Axis Combined Print Strategy Demo SampleThe 5-axis samples were not able to withstand as much compressive force as the 3-axis horizontal samples. This would indicate that only part of these samples acted in compression until failure.

This would indicate that only part of these samples acted in compression until failure. The rest of the sample would then be acting in tension, to create a bending moment in the sample as illustrated in Figure 10.Applied Loading◊Figure 10: Illustration of Internal Compression/Tension Present in 5-Axis SamplesWhile this mechanical behaviour did not allow the 5-axis samples to withstand as much force as some of the 3-axis samples, it does indicate that these samples behave more isotopically, since all three build orientations had rather similar mechanical properties. This is an interesting result that could lead to further research and applications in industry.1827ConclusionResearch into 5-axis ME is currently rather underdeveloped. As such, experimentation is required to improve the understanding of the limitations and benefits of this technology. The 3-axis ME results produced by this work did not generate the expected failure behaviour.

The 3-axis ME results produced by this work did not generate the expected failure behaviour. However, the 5-axis samples did display the predicted mechanical behaviour in terms of which samples performed best in the compression loading test. This indicates that a combination print stratergy has some potential to improve mechanical behaviour in ME samples, but that further research is required to explore its applicability. The anisotropy of most current ME process was shown most dramatically in the difference between the 3-axis vertical samples and the 3-axis horizontal samples. Orientating the print through 90° changed the load-bearing capability of samples by a factor of 3.5. Overall, 5-axis deposition along principle stress vector directions enabled the anisotropic nature of ME to be reduced.

Overall, 5-axis deposition along principle stress vector directions enabled the anisotropic nature of ME to be reduced. Such behaviour has the potential to expand the use of this manufacturing process for mechanical prototypes and end-use products.A limitation of the 5-axis ME process that was used was that there is no automated software available to create the tool paths required for printing. Generating bespoke tool paths for samples could lead to further research in what is the best way to align material in a 5-axis context.A future area of research will be to continue to explore 5-axis ME with the inclusion of reinforcing fibres that can be aligned in the direction of printing. At the time of writing this work is currently being carried out with a view to further publications.References 1. Wohlers T, Gornet T. History of Additive Manufacturing. Wohlers Rep 2014 - 3DPrint Addit Manuf State Ind. 2014;1–34.2. Wohlers TT, Associates W.

viewed PaperCOMPARATIVE STUDY OF MECHANICAL PROPERTIES OF ALUMINUM ALLOY A356 (AL-12SI) FABRICATED BY DIRECTED ENERGY DEPOSITION METHOD AND PRESSURE DIE CASTINGRagavanantham Shanmugama, Jishu Chandranb, Mohanavel Vinayagamc, OsamaFakrondSeth Dennisone, Seth Rominefa,d,e,f School of Engineering, Mathematics and Technology, Navajo Technical University, NM- 87312, USAb Department of Mechanical Engineering, Vel Tech Rangarajan Dr.Sagunthala R&D Institute of Science and Technology, Chennai, IndiacDepartment of Mechanical Engineering, Bharath Institute of Highere Education & Research, Selayur, IndiaAbstract:-Additive Manufacturing technology is rapidly gaining traction in many manufacturing applications due to its process parameters control and wide range of applications. There are other AM technologies accessible, however Direct energy deposition (DED) is a critical approach in metal matrix additive manufacturing.

There are other AM technologies accessible, however Direct energy deposition (DED) is a critical approach in metal matrix additive manufacturing. The significant two mechanical properties tensile strength and hardness of additively created Al-12Si Aluminum alloy by DED process and pressure die casted Al-12Si components are compared in this study. The strength and modulus of the DED and PDC manufactured components were identical if the load direction in the UTM machine was the same as the construction directions, however other mechanical parameters differed slightly. Mechanical qualities of fabricated products made from reused powders were also comparable to those made from unused powder.Keywords: Direct energy deposition (DED), AlSi12 alloy, Pressure die casting, ultimate tensile strength, hardness, building directions1.

INTRODUCTIONAdditive manufacturing (AM) technology provides significant advantages over traditional production procedures against components such as casting moulds, punches, and dies. This method uses a computer-generated 3D model and prepares the data for printing the specified object. Unlike traditional manufacturing procedures, complex pieces with acceptable accuracy can be manufactured in one step, considerably reducing the number of parts assembled in a single product [1]. Furthermore, bespoke products can be created without sacrificing inventory spare or production lead time. Because of its various advantages over previous manufacturing techniques, AM is regarded as a unique method for developing and producing customised components for the space bus, automotive, and biomedical industries [2-3]. Previously, the fuel injector nozzle was assembled from numerous pieces in the aircraft industry.

Previously, the fuel injector nozzle was assembled from numerous pieces in the aircraft industry. In addition introduction of AM technology, it is now possible to create a lightweight designed nozzle at a substantially lower cost. It is simple to create patient-friendly equipments utilising AM processes by using patient medical photos, and AM processes may also produce high-performance biocompatibility devices. Burner tips that mix and swirl in complicated forms composed of high-temperature materials to save cost, fuel, Prototyping, quick1838fabrication, and service of industrial equipment such as punches, dies, bespoke tooling, and so on, are made more accessible in the automotive industry [4-6]. Significant advancements in the underlying concepts of AM metal processing, software and hardware, and feedstock technology have allowed AM to become a cutting-edge technology over the previous two decades.

It has now achieved an acceptable level as a result of strong development in commercial system sales. Initially developed in research facilities, AM technology is now being illustrated and accepted by industry [7].Mechanical anisotropy and internal stress are two areas where AM and traditionally produced components differ. Unique faults to the AM process must also be carefully communicated in aerospace fields where components such as fuel injector nozzles, etc., are subjected to more fatigue [8]. other components, such as turbine blades, are still in the early stages of development. Powder metallurgy, beam welding, and prototyping are the fundamental concepts for developing additive alloy production. Nonetheless, available knowledge from these methodologies does not demonstrate many important aspects of additive manufacturing. A rather mature expertise in welding and powder metallurgy has been created through many decades of research effort.

A rather mature expertise in welding and powder metallurgy has been created through many decades of research effort. This has paved the way for future study and development in metal alloy additive fabrication [9]. Still, it appears to be a lengthy and complicated path.Aside from that, AM technology offers some distinct advantages that are propelling this sector forward, notably in metallic materials.Aluminium is widely used in die-casting, although it is prone to breaking or shrinking at high temperatures, thus it is frequently alloyed with copper or silicon. The addition of these elements to aluminium increases its hardness and fluidity. Aluminium is used to make components with thin walls and complex shapes because of its great dimensional stability. Aluminium's corrosion resistance makes it useful in components that are subjected to thermal or electrical energy exposure [10].

Aluminium's corrosion resistance makes it useful in components that are subjected to thermal or electrical energy exposure [10]. The compositions of AlSi10Mg and AlSi12 are close to the eutectic points, which have a lower melting point and a narrower solidification temperature range. As a result, general alloys have good casting characteristics and lower shrinkage porosity, making them better for DED processing. The maximum alloys SLM research is focused on processing [11]. The alloy, on the other hand, has recently gotten a lot of interest reason it's a near-eutectic alloy with good castability and strength [12]. Heat treatment of DED-processed AlSi12 specimens was thoroughly investigated. Heat treatment has been discovered to modify the microstructure and effect mechanical characteristics [13]. Furthermore, it have found the protective gas employed during the DED process has no effect on the AlSi12 densification behavior. They do, however, have a minor impact on ductility.

Furthermore, it have found the protective gas employed during the DED process has no effect on the AlSi12 densification behavior. They do, however, have a minor impact on ductility. The cast alloy was compared to AlSi12 specimens generated by DED (as-fabricated and heat- treated) in terms of tribological and corrosion behaviour. The effect of cooling rate on microstructure and characteristics has been determined [14]. The table.1 shows AlSi12 alloy composistionsChemical composition, wt. % Si Al Cu Mg 11.2 0.45 0.42 0.12 0.10 0.10 0.10 0.10 0.08 balFe Mn ZnTiPbNiTable 1. chemical composition of raw AlSi12 alloy2. AM PROCESSES1839The AM technique is also called as 3D printing, There are many different types of additive manufacturing (AM) procedures, each with its own way of layer manufacturing, material, and machine technology. The ASTM developed standards that divide Additive Manufacturing into seven distinct categories.A brief summary of each technique is provided.

The ASTM developed standards that divide Additive Manufacturing into seven distinct categories.A brief summary of each technique is provided. Using UV laser light, vat polymerization is used to create a product or model from a pool of liquid polymer. UV light is utilized to make solid the polymer liquid at a specific area based on computer-generated design data. Once each layer of the thing being made has formed, the platform lowers it. The support structure must be introduced throughout the product's construction; after the product is finished, the support structure must be removed in order to obtain the final product.Powder and a binder are the main components of binder jetting. The binder acts as a glue to hold the powder layers together. Liquid binder and powdered building material are both common. Construction and binding materials are deposited in layers as the print head translates horizontally along the x and y axes.

Liquid binder and powdered building material are both common. Construction and binding materials are deposited in layers as the print head translates horizontally along the x and y axes. The object is deposited onto its building platform after each layer. Because of the binding technique, the material properties are manufactured suitable for structural components. No matter how fast the printing process is, further post-processing might take a significant amount of time. In the same way as previous powder-based manufacturing methods, the printed object is self-supported inside the powder bed and removed from the unbound powder after finishing. In the material extrusion process material is pushed through a nozzle where it is heated, and then deposited layer by layer. The nozzle and platform may move horizontally and vertically after each layer is deposited.

The nozzle and platform may move horizontally and vertically after each layer is deposited. Many low-cost, home-based 3D printers uses this technology.There are additional additive manufacturing methods namely ultrasonic additive manufacturing and laminated object manufacturing. Metal sheets and ribbons are used in ultrasonic additive manufacturing but laminated object manufacturing make use of cross hatching technique for smooth removal of printed parts after the process. The printing parts made by laminated objects manufacturing are not structurly sound however they often used for visual modelling. In ultrasonic additivie manufacturing commonly used materials are Aluminium, copper, stainless steel, and titanium, after printing the parts CNC machining required for removing unbound metal, which is commonly done during the welding process. PBF processes employ a laser or an electron beam to melt and fuse the powder.

PBF processes employ a laser or an electron beam to melt and fuse the powder. Although electron beam melting (EBM) needs a vacuum, it may be used to create useful components out of metals and alloys. In all PBF methods, the powder material is disseminated across the previous layers. This may be done using a number of tools, such as a roller or a blade. A hopper or reservoir underneath or to the side of the bed supplies fresh material.3. EXPERIMENTAL PROCEDUREIn this present study laser engineering net shaping method was used to manufacture three dimensional specimens. Samples are fabricated by using LENS™, Optomec Inc, the model of the specimen is designed by using solid works software, then the model is converted into STL file before feed into system. Various parameters used to manufacture the specimens are indicated in Table 2. AlSi12 alloy powder was prepared by gas atomization process wherein diameter of powder particles is kept in the range between 5µm to 15µm.

AlSi12 alloy powder was prepared by gas atomization process wherein diameter of powder particles is kept in the range between 5µm to 15µm. By using LENS various samples are prepared for doing mechanical test, a Universal testing machine [model 150ST, Transverse speed: .05 to 500 mm/min] with a maximum capacity of 150 kN is used. Rockwell hardness tester is used to find out hardness with B scale used to check it out. Al-9Si1840and Al-12Si two basic alloy groups used in high pressure die casting. Only a very tiny fraction of the whole market is made up of other alloys. A force is imparted to the dies in use in pressure die casting. Based on the specimen, the temperature of the dies is kept between 120°C and 280°C. An alloy metal shot sleeve is filled with 600–750°C alloy metal. The plunger pressurises the molten metal, which then goes forward and into the die.

An alloy metal shot sleeve is filled with 600–750°C alloy metal. The plunger pressurises the molten metal, which then goes forward and into the die. As the molten metal travels through the die's runners, it passes through a tiny channel known as a "gate" before entering the casting's hollow. In aluminium die casting, high pressures of 30–150 MPa are used, as well as the molten metal is reduced in the die to the point where it freezes. After that, the ejector pins are used to extract the cast item from the dies. To cool and lubricate the dies, the water/oil emulsion is subsequently applied. Afterward, the dies are reunited and the procedure is performed once again [16-18].4. RESULT AND DISCUSSION4.1. Tensile propertiesTensile test was performed aluminium alloy specimen manufactured by direct energy deposition and pressure die casting based on ASTM E8M standard.

RESULT AND DISCUSSION4.1. Tensile propertiesTensile test was performed aluminium alloy specimen manufactured by direct energy deposition and pressure die casting based on ASTM E8M standard. Three samples are taken for test, below table and graph indicate mean variation in three mechanical properties of ultimate tensile strength, percentage of elongation, yield strength. The result variation shows that additively manufactured AlSi12 alloy have commendable mechanical properties over PDC manufactured aluminium alloy specimen.This superiority in mechanical strength is due to fine silicon particles uniformaly distributed over the each layers in direct energy deposition process.Fractured specimen exihibit finely dispersed particle as ductile behavior prevails.Sr. no.1 2specimenAM- DED PDCStrength (MPa) 420.19 165.08% Elongation4.21 1.10Yield Strength (MPa) 315.20 132.35Table 2.

no.1 2specimenAM- DED PDCStrength (MPa) 420.19 165.08% Elongation4.21 1.10Yield Strength (MPa) 315.20 132.35Table 2. Variations of mechanical properties500400300AM- DED200PDC1000UTS(MPa)Yield Strength (MPa)Fig 1. Comparison of UTS and yield strength of two specimen18414.543.532.521.5AM- DED1PDC0.50AM- DEDPDC% of ElongationFig 2. Comparison of % of elongation of two specimen4.2 Hardness testIn most of the mechanical properties, hardness is well known and basic . In this experiment hardness is checked on rockwell hardness tsting machine. To test three separate areas of the top surface, both specimens are made using two completely different methods. Table 3 shows the average hardness value. The hardness value was determined using the B scale. A steel ball indenter with a resolution of 1/1600 was employed with a 150 kg main load.

Table 3 shows the average hardness value. The hardness value was determined using the B scale. A steel ball indenter with a resolution of 1/1600 was employed with a 150 kg main load. Due to the tiny grain microstructure created during the Permit and the quick powder solidification, the DED samples have a greater hardness than cast alloy, as shown by the results of the Hardness Test. The hardness (RB61) of the AM sample is much higher than that of the pressure die cast sample (RB19). Using these findings, it may be concluded that additively created specimens have the correct particle dispersion. As predicted, the hardness values match the strength results.Sr. no. 1Sample AM- DEDHardness (RB) 612PDC19Table 3. Variations of mechanical properties5. CONCLUSIONSpecimens manufactured by direct energy deposition and pressure die casting were investigated under various mechanical tests.

Variations of mechanical properties5. CONCLUSIONSpecimens manufactured by direct energy deposition and pressure die casting were investigated under various mechanical tests. The specimens shows high level of variations in mechanical properties such as tensile strength, percentage of elongation and hardness. AM samples performed betted by all metrics: higher yield strength, higher ultimate tensile strength, more elongation at rupture, and higher hardness. Pressure die casting specimen, bonding between the particles are much lower than additively manufactured specimen. AM- DED specimen shows significantly improved mechanical properties due to inter layer bonding which results in better tensile strength, also fine grain structure, grain size and better cooling rate. In addition to residual stress, metallurgical flaws such as pores, fractures and surface roughness is a greater problem in additive manufacturing components.

In addition to residual stress, metallurgical flaws such as pores, fractures and surface roughness is a greater problem in additive manufacturing components. However, the underlying mechanisms for the creation of these abnormalities remain a mystery.1842REFERENCES[1] Milewski JO. Additive manufacturing of metals. Springer series in materials science, vol. 258. Springer; 2017[2] Zhang, J.L.; Song, B.; Wei, Q.S.; Bourell, D.; Shi, Y.S. A review of selective laser melting of aluminum alloys: processing, microstructure, property and developing trends. J. Mater. Sci. Technol. 2019, 35, 270–284.[3]Shanmugam, R., Ramoni, M., Chandran, J., Mohanavel, V. and Pugazhendhi, L., 2021. A Review on the significant classification of Additive Manufacturing. Journal of Physics: Conference Series, 2027(1), p.012026.[4] Kruth, J.P.; Levy, G.; Klocke, F.; Childs, T.H. Consolidation phenomena in laser and powder-bed based layered manufacturing. CIRP Ann.

COMPARATIVE STUDY OF METALLIC ADDITIVE MANUFACTURING POWER CONSUMPTION M. Baumers*, C. Tuck*, R. Hague*, I. Ashcroft* and R. Wildman* *Additive Manufacturing Research Group, Wolfson School of Mechanical and Manufacturing Engineering, Loughborough University, LE11 3TU, UKAbstractEfficient resource utilisation is seen as one of the advantages of Additive Manufacturing (AM). This paper presents a comparative assessment of electricity consumption of two major metallic AM processes, selective laser melting and electron beam melting. The experiments performed for this study are based on the production of a common power monitoring geometry. Due to the technology’s parallel nature, the degree of build volume utilization will affect any power consumption metric. Therefore, this work explores energy consumption on the basis of whole builds - while compensating for discrepancies in packing efficiency.

Therefore, this work explores energy consumption on the basis of whole builds - while compensating for discrepancies in packing efficiency. This provides insight not only into absolute levels of power consumption but also on comparative process efficiency.IntroductionThe potential impact of global warming may make a fundamental change in economic activity necessary. For the UK alone, recent figures estimate emissions of the six most important greenhouse gases at around 628.3 million tonnes of carbon dioxide equivalent (MtCO2e) per year (DECC, 2010). The most significant source of greenhouse gas emissions in the UK is energy generation, creating around 219.7 MtCO2e per annum. Emissions resulting directly from industrial processes contribute 16.7 MtCO2e.A potential reduction of the carbon footprint of manufacturing through the adoption of Additive Manufacturing is the subject of the ATKINS project conducted at Loughborough University.

The authors report system energy consumption in terms of an energy consumption rate, measuring kWh consumed per kg of part geometry. Tab. 1 shows the range of energy consumption rates observed by Luo et al. for the three AM variants:SLALSFDM20.70 – 41.38 kWh/kg29.83 – 40.09 kWh/kg23.08 – 346.4 kWh/kgTable 1: Ranges of energy consumption rates, Luo et al., 1999In a similar framework, Sreenivasan and Bourell (2009) study power consumption of LS. Apart from reporting the energy consumption rate (14.5 kWh/kg), the authors also citemean power consumption of the investigated LS system, a 3D Systems Vanguard HS + HiQ, measured at 19.6 kWh over the investigated build. It is noteworthy that there is little research on the energy usage of the metallic variants of AM. In a cross-platform study of the effect of build parameters (such as part orientation) on AM energy consumption, Mognol et al.

In a cross-platform study of the effect of build parameters (such as part orientation) on AM energy consumption, Mognol et al. (2006) present power consumption estimates for the 3D Systems Thermojet 3D printer, the Stratasys FDM 3000 machine and the EOS M 250 direct metal laser sintering (DMLS) system. Tab.2 shows mean power consumption in the system’s idle and busy states, as well as the minimum energy consumed to manufacture a power monitoring geometry, as reported by Mognol et al.:3D PrintingFDMDMLSIdle mean power consumption Busy mean power consumption Energy consumed per part (min.)0.69 kW0.88 kW2.1 kWh0.53 kW0.57 kW0.5 kWh2.00 kW4.00 kW32 kWhTable 2: mean power and energy consumption, Mognol et al., 1999These figures appear to show that the DMLS process (and perhaps metallic AM in general) consumes disproportionately more energy than the non-metallic technology variants.

Nevertheless, the adoption of metallic AM may result in a significant energy saving when compared to conventional manufacturing techniques and their products. Further energy savings may arise during the product life cycle of AM produced goods (see ATKINS feasibility study, 2008). These may be due to lower raw material usage, more efficient distribution methods, better part performance and improved recycling.The goal of this study is to provide generally applicable information on electric energy consumption of modern metallic AM platforms. Power monitoring experiments were performed on an MTT SLM250 selective laser melting (SLM) system and an Arcam A1 electron beam melting (EBM) machine. A particular emphasis was placed on designing experiments and summary metrics that treat AM as a parallel manufacturing technology, allowing the contemporaneous production of different parts, an aspect missing from the discussed studies of AM energy efficiency.

Previous models of AM input efficiency, albeit of financial inputs, describe efficient (i.e. minimum cost) technology usage only if they are performed on the basis of fully packed build volumes (Hopkinson and Dickens, 2003; Ruffo et al., 2006). This research demonstrates that the same is true in the assessment of energy inputs.Experimental setupFor this research, a series of power monitoring experiments was conducted using a portable power monitoring setup featuring Yokogawa’s CW240 digital multi-purpose power meter (for specifications, see the Yokogawa product guide, 2007). For the Arcam A1 system, power consumption was measured in a three-phase-four-wire configuration (400 V), with a total of eight probes (four current clamps and four voltage probes) connected to the system’s multi phase power supply (Fig.1).

However, the meter is also capable of measuring balanced and unbalanced loads in three-phase-three-wire-three-current connections.The newest version of MTT’s SLM250 operates on a single phase 240 V power supply. Hence, the power meter was set to a one-phase-two-wire configuration, with one current clamp and one voltage probe attached to both the phase and the neutral, totalling four probes.Voltage probes 3 phases, 1 neutralCurrent clamps (type 96033) 3 phases, 1 neutralFigure 1: Three-phase-four-wire configuration, artwork taken from: User’s Manual IM CW240E, Yokogawa, 2004In the context of this study, the main variables of interest are mean real power consumption per measurement cycle (measured in W and denoted internally by the meter as ‘P_AVE(W)_1’), and total cumulative energy consumed (measured in Wh and denoted as ‘Wh+_INTEG(Wh)_1’).

The power meter was configured to a 1s measurement cycle, this cycle length being the shortest for which all 137 measurement variables are available. Two power monitoring experiments were carried out for each AM platform:1. In the first experiment the production of a full build volume of parts is studied, giving data on energy consumption when the machine is operating at full capacity.2. The second experiment surveys the production of a single part located in the centre of the work space. These data were required to perform a normalisation of packing density and to analyse efficiency gains from multi-part production.To measure total process energy consumption, the AM machines were monitored during the entire build time.

Fixed process steps preceding the actual build, for example, energy expended for bed heating or vacuum drawing, are included.An understanding of the relationships between part geometry and energy usage may be won by using a standardised power monitoring geometry. The layer-wise operating principle of the studied AM systems provides an opportunity to design a common geometry for use in such experiments. By varying the cross-sectional shape along the part’s Z-axis, the effect of geometric parameters on energy consumption can be isolated. The performed power monitoring experiments use the part shown in Fig.2. This part features variation in two parameters, geometric/topological complexity and part volume.Z-orientation (part height: 24 mm)75 mm75 mmFigure 2: The standardised power monitoring geometryIn the part’s lower half, the effect of geometric / topological complexity on power consumption can be assessed.

This is done by varying a simple measure associated with such complexity, the ratio of cross section perimeter P and area A, along the Z axis. This also lends the part a ‘spider’ shaped footprint with a cavity in the centre (not visible in Fig.2). Moving up vertically, the part cross-section changes into a simple square without cavity, at a Z-height of 12 mm.In the top half of the geometry (>12 mm Z-height), the effect of a reduction of cross section area can be studied. This is achieved by shrinking area A along the Z-axis down to a value of zero - creating a pyramidal tip. The described parameter variation can be expressed graphically, as done in Fig.3. Part cross section (solid graph) is kept relatively constant up to a Z-height of 14 mm, above which it is gradually reduced to zero. Geometric complexity (dotted graph) is reduced in the lower half of the geometry.

Geometric complexity (dotted graph) is reduced in the lower half of the geometry. However, this simple measure starts rising above 14 mm Z-height, due to part area A diminishing faster than the perimeter P – however, the square shape of the cross section remains, and with it the real level of complexity.Figure 3: Variation of geometric parameters in the power monitoring geometryWhile the current paper does not report the effects of geometric parameters on energy consumption, a further reason to choose the ‘spider’ shape was that it has a relatively large footprint in the X / Y dimensions. It thereby limits overall packing density, making the power monitoring builds faster and more economical. This is particularly effective for AM approaches that require every part to be attached to the build platform (referred to as 2½ dimensional build volume packing).

This is particularly effective for AM approaches that require every part to be attached to the build platform (referred to as 2½ dimensional build volume packing). A further consideration in the design of the part is the limitation of negative wall angles (<45º) to ensure that the investigated processes do not require supports for overhangs. Nevertheless, on some systems (such as the MTT SLM250) support structures may be required to attach the parts to a base plate.When analysing energy consumption on the basis of full builds of identical parts, it can be expected that discrepancies in packing density arise across AM platforms, due to different build volume sizes and shapes. This makes an adjustment of the observed power consumption data necessary, which is achieved by normalising packing density.

This makes an adjustment of the observed power consumption data necessary, which is achieved by normalising packing density. Making the simplifying assumption that for each AM system the relationship between build energy consumption and the number of parts in the build volume is linear, packing density can then be adjusted by extrapolating from the two experimental results obtained for each AM platform (full build and single part). The first step is to adjust packing density, the number of parts n parts per area A of the build platform, according to some standardised density (nSTANDARD/ASTANDARD):AnAnnnSTANDARDSTANDARDADJUSTEDIdeally, the standard packing density would be the highest observed density in the experiments. In the sample of AM systems assessed for this paper, the Arcam A1 EBM machine featured the highest density.

In the sample of AM systems assessed for this paper, the Arcam A1 EBM machine featured the highest density. It should be noted that this is not indicative of system efficiency in any way; it merely shows how well the power monitoring geometries fit into a particular build volume. A corrected measure for total energy consumption during the full build experiment (EADJUSTED) can then be calculated using nADJUSTED in conjunction with the simple intercept theorem; EFULL BUILD and ESINGLE PART denote the total energy consumed during the full build and the single part experiments:PARTSINGLEBUILDFULLPARTSINGLEBUILDFULLADJUSTEDADJUSTEDEnEEnE1)()1(To obtain the adjusted measure for the energy invested per part (EPER PART, ADJUSTED),(2) is divided by nADJUSTED:ADJUSTEDADJUSTEDADJUSTEDPARTPERnbuildperconsumedEnergyE,AM platforms investigatedThe two studied variants of metallic AM share numerous features.

Both operate by adding part material layer by layer in a powder bed. Nominal build volume sizes are also relatively similar, with 250 * 250 *300 mm on the MTT SLM250 and 200 * 200 * 180 mm(1)(2)(3)on the Arcam A1. Moreover, both processes require that the manufactured parts are attached to a removable base plate. During operation, this plate is lowered by a small increment after each layer has been completed. However, the machines do exhibit significant differences. Tab.

While the SLM system scans the surface with a 200 W fibre laser, transferring energy in the form of light, the EBM system melts the metal powder with an electron beam (maximum output: 3000 W). According to Strutt (1980), energy transfer efficiency by electron beam can be around 10 times greater than for certain laser types. However, this value was obtained from a comparison between electron beams and CO2 lasers; the SLM machine employs a more efficient fibre laser operating at 1.06 μm wavelength. Thus, the comparatively lower efficiency of laser-based systems may be due to melt pool reflectivity. A drawback of using an electron beam to scan the powder bed is that this method imposes restrictions on the usable build materials, magnetic metal powders cannot be used.The machines’ process atmospheres also differ strongly. The build activity in the Arcam A1 is performed in a near vacuum (with the addition of He), drawn by a powerful turbo pump before the scanning begins.

The build activity in the Arcam A1 is performed in a near vacuum (with the addition of He), drawn by a powerful turbo pump before the scanning begins. This allows a hot process, with bed temperatures of around 700º C. The high power of the electron beam, together with its efficient energy transfer, makes additional bed heating elements unnecessary – the powder bed is heated by evenly scanning the bed surface with the electron beam. In contrast to this, the MTT SLM250operates with an Ar build atmosphere at just over normal pressure (~ 15 mbar over outside atmosphere). The process belongs to the class of ‘cold’ AM processes with bed temperatures between 100º C and 300º C. While this machine possesses build chamber heating elements, this function was deactivated during the experiments in order to run the machine at maximum energy efficiency. The MTT SLM250’s external cooling system draws power independently and had to be monitored using a second power meter.

The MTT SLM250’s external cooling system draws power independently and had to be monitored using a second power meter. Apart from the presented measures of real power consumption, external chiller energy consumption was included in all other results presented in this paper.In order to be reflective of machine usage in established applications, the experiments were performed with different material grades and layer thicknesses on the two machines. It was expected that the low layer thickness (50 μ) on the MTT SLM250 will result in lower process speed and higher energy consumption, due to the greater number of layers required to build the standardised power monitoring geometry.

It should be noted that the advantages of small layer thicknesses, such as higher geometric tolerances and diminished surface quality, are not factored into the results presented.Results and DiscussionTwo build experiments were performed on both AM platforms, one full build experiment containing as many power monitoring geometries as possible and one build with a single part in the centre of the build platform. As expected, the packing densities for the full build volumes vary: while the build envelope of the MTT SLM250 is able to hold 6 parts, the Arcam A1 is limited to 5 power monitoring geometries. To give a concise overview, the results of the power monitoring experiments are presented in Tab. 4:SLMEBMMachine typeMTT SLM 250Arcam A1Number of parts in full buildFull buildBuild timeSingle partMean real power consumedFull buildSingle part6 1519 min (excl. supports: 1153 min) 351 min (excl. supports: 286 min) 1.10 kW (excl. chiller)0.92 kW (excl.

supports: 1153 min) 351 min (excl. supports: 286 min) 1.10 kW (excl. chiller)0.92 kW (excl. chiller)5304 min192 min2.22 kW2.01 kWFull build44.26 kWh11.04 kWhTotal energy consumedSummary metrics (adjusted)Single part Parallel manufacturing factor Energy consumed per part Energy consumed per g (assuming 100% dense parts) Energy consumed per cm³9.14 kWh0.817.34 kWh0.031 kWh0.249 kWh6.41 kWh0.342.21 kWh0.017 kWh0.075 kWhTable 4: Power monitoring resultsThe table presents results on five different categories: part quantity, build time, real power consumption, total energy consumed and summary metrics. Of the surveyed machines, the Arcam A1 was the fastest to complete the full build experiment in 304 min, compared to 1519 min for the MTT SLM250. Subtracting the time used for the creation of the supports, the build was completed in 1153 min.

Subtracting the time used for the creation of the supports, the build was completed in 1153 min. The higher speed of the Arcam A1 can partially be explained by the fact that it has a smaller build volume than the MTT SLM250. Moreover, it operates with an effective electron beam scanning system without moving parts, capable of high scan speeds. Another factor of particular importance is that layer thickness on the Arcam A1 (80 μ) is 1.6 times larger than on the MTT SLM250 (50 μ). It should also be noted that the supports connecting the parts to the platform were not optimised for the MTT SLM250, resulting in a further slowdown of build speed.The measured mean real power consumption describes the average amount of AC power the machine is consuming over the course of the build, including any fixed heating up, atmosphere generation and cooling down procedures. The power supply specification is chosen by the manufacturers based on this characteristic.

The power supply specification is chosen by the manufacturers based on this characteristic. However, this measure is not very helpful in the analysis of AM energy efficiency as it omits process time and productivity. Nevertheless, existing studies on AM energy efficiency report this result (Mognol et al., 2006; Sreenivasan and Bourell, 2009). The values observed during the current research appear significantly lower than the results presented in the literature. Mognol et al. report an average build time power consumption of 4.00 kW for the (now obsolete) EOS M 250 DMLS system, this appears high compared to an energy consumption of 1.10 kW for the MTT SLM250, which is technologically similar.Total energy consumption is a more applicable measure of true process efficiency, in particular when recorded for a full build. In the performed experiments, both in the full build and in the single part experiments, the Arcam A1 showed the lowest total energy expenditure (11.04 and 6.41 kWh).

In the performed experiments, both in the full build and in the single part experiments, the Arcam A1 showed the lowest total energy expenditure (11.04 and 6.41 kWh). As discussed above, this is not surprising considering the smaller build volume and the high layer thickness. Furthermore, EBM uses titanium (Ti-6Al-4V) as raw material. Titanium has a lower density than steel, combined with a roughly similar specific heat capacity (~ 0.5 kJ/kg K). It should be noted that titanium has a higher melt point and a different energy absorption rate than steel. Measuring total energy consumption allows the the calculation of a statistic reflecting contemporaneous production of multiple parts, labelled the ‘parallel manufacturing factor’:the energy savings availablethroughPARTSINGLEBUILDFULLEnEfactoringmanufacturParallel/(4)This factor gives an indication of the effect on energy consumption of parallel production of multiple parts.

It shows how well the energy investment associated with the process can be amortised over multiple parts included in each build. A low value indicates that the savings available from parallel production are large: the factor of 0.34 recorded for the Arcam A1 implies that if multiple parts are manufactured per build instead of one, only 34 % of the energy needs to be invested per part. With a parallel manufacturing factor of 0.81, the MTT SLM250 offers lower energy savings of this kind. A more powerful laser, faster scanning and a larger build volume could improve this value, however.Appreciating the discussed impact of packing density, summary metrics (based on full build energy consumption) are reported in adjusted form. Information on the following three metrics is reported: energy consumption per part, energy consumed per g and energy consumed per cm3 of part volume.

Information on the following three metrics is reported: energy consumption per part, energy consumed per g and energy consumed per cm3 of part volume. In terms of energy consumption per part, the results show that the SLM process consumes around 7.34 kWh while the EBM process uses around 2.21 kWh. Keeping in mind that differences in material, layer thickness and surface quality may account for this, it appears that the Arcam A1 consumes around 70% less energy per part than the competing SLM process. This is surprising as mean real power consumption of EBM is much higher (2.22 kW vs. 1.10 kW without chiller). In terms of overall energy consumption, though, this is more than offset by a higher process speed (304 min vs. 1519 min). The measure for energy consumed per g of part volume is based on fully dense parts, assuming a density of 8.00 g / cm³ for the parts produced in stainless steel grade SAE 316L on MTT’s SLM250 and 4.43 g / cm³ for the Ti-6Al-4V parts from the Arcam A1.

This summary metric can be used to mimic the energy consumption rates of AM systems in the literature, as reported for polymeric LS by Luo et al. (1999) and Sreenivasan and Bourell (2009):Lou et al. (1999)Sreenivasan and Bourell (2009)This researchSystem typeProcess classDTM Sinterstation 2000, Sinterstation 2500 LS3DS Vanguard HiQ+HSLSMTT SLM250SLMArcam A1EBMMaterialPolyamidePolyamideSAE 316LTi-6Al-4VEnergy consumption rate40.09 kWh / kg , 29.83 kWh / kg14.5 kWh / kg31 kWh / kg17 kWh / kgTable 5: Energy consumption rates for AM systemsEven though the processes analysed for this research generate metal parts and the previous work focuses on polymeric AM, the obtained results roughly fall in line with the published energy consumption levels for LS.

However, the discrepancy between this result and the observation that metallic AM consumes more energy than polymeric AM (Mognol et al., 2006) suggests that more research is needed.ConclusionsThe comparative assessment of electricity consumption of the SLM and EBM processes demonstrates that studies on input efficiency for parallel manufacturing systems can lead to different results if full build volumes are considered instead of one-off parts. This will be particularly relevant for systems exhibiting a low parallel manufacturing factor, as defined above (4). The single part build experiment on the Arcam A1 (factor 0.39) resulted in an energy investment of 6.41 kWh, compared to an adjusted full build energy consumption per part of 2.21 kWh.

The single part build experiment on the Arcam A1 (factor 0.39) resulted in an energy investment of 6.41 kWh, compared to an adjusted full build energy consumption per part of 2.21 kWh. For the MTT SLM250 (factor 0.81), the picture is similar but not as pronounced: the single build required in an energy investment of 9.14 kWh, compared to an adjusted full build energy consumption per part of 7.34 kWh.It has also been demonstrated that a standardised testing part can lead to meaningful results in full build studies of AM energy consumption. An important step is, however, tocompensate for differences in packing density. This ensures that results that are not distorted by indivisibilities resulting from fixed part size.The lower energy consumption measured for the EBM platforms appears to be, at least partially, due to differences in material and layer thickness. While the specific heat capacities of stainless steel and titanium are similar, the density of titanium is much lower.

While the specific heat capacities of stainless steel and titanium are similar, the density of titanium is much lower. Hence, to heat a volume measure by 1º C, steel requires more energy. Moreover, the presented metrics do not reflect the advantages of a low layer thickness - such as improved surface finish and tighter dimensional tolerances. For future studies, it would be advisable to compensate for variations in layer thickness. Furthermore, it should be noted that this study has not dealt with part supports in a systematic way. Unlike the experiments performed on the Arcam A1, the experiments performed on the MTT SLM250 included substantial support structures connecting the parts to the base plate (created using the Marcam AutoFab package).In the context of the ATKINS project, it will be important to substantiate claims that AM is a more energy efficient alternative to conventional manufacturing.

This can be done by comparing summary metrics, such as kWh/cm³ or kWh/g, to corresponding data collected for conventional manufacturing processesAcknowledgementsThe authors would like to use this opportunity to express their gratitude to the following individuals who have expertly performed the build experiments. The experiments took place at several locations: at the AMRG at Loughborough University (UK), thanks go to Mark Hardy; at Arcam AB (Gothenburg, Sweden), the authors thank Dr. Ulric Ljungblad, Isak Elfström and Daniel Jonasson; and at MTT Technologies Group (Stone, UK) where the authors thank Dr. Chris Sutcliffe, Kevin Hart and Paul Smith for their support.ReferencesArcam AB, 2010. Arcam A1 – The future in implant manufacturing. [Online]. Available from: http://www.arcam.com/CommonResources/Files/www.arcam.com/Documents/Produc ts/Arcam-A1.pdf [Accessed: 15.06.2010]ATKINS feasibility study. Downloads, 2008. [Online].

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA Comparative Study on the Microstructure and Texture Evolution of L-PBF and LP-DED 17-4 PH Stainless Steel during Heat TreatmentP.D. Nezhadfar1,2, Paul R. Gradl3, Shuai Shao1,2, Nima Shamsaei1,21Department of Mechanical Engineering, Auburn University, Auburn, AL, USA 2National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL, USA 3 NASA Marshall Space Flight Center, Propulsion Department, Huntsville, AL 35812, USACorresponding author: Email: shamsaei@auburn.edu Phone: (334) 844 4839AbstractThis study aims to characterize the microstructure and crystallographic texture of 17-4 PH stainless steel (SS) manufactured with laser powder directed energy deposition (LP-DED) and laser powder bed fusion (L-PBF), in both non-heat treated and heat treated conditions.

It is found that the non-heat treated LP-DED 17-4 PH SS possesses coarse columnar ferrite grains decorated with Widmanstätten ferrite grains, whereas the L-PBF counterpart has very fine and mostly equiaxed ferrite grains along with lath martensite. An identical stress relief (SR) temperature is obtained for both the L-PBF and LP-DED 17-4 PH SS samples based on the phase diagrams generated using Thermo-Calc. software. The SR step prior to CA-H1025 heat treatment resulted in texture weakening and slightly refined the grain structure.

software. The SR step prior to CA-H1025 heat treatment resulted in texture weakening and slightly refined the grain structure. The non-heat treated L-PBF 17-4 PH SS sample possesses strong cube and γ-fiber textures, while the texture transfers to weaker γ-fiber components after performing SR-CA-H1025 heat treatment.Keywords: Laser powder directed energy deposition (LP-DED); Laser powder bed fusion (L- PBF); Stainless steel; Microstructure; TextureA substantially improved version of this paper appears in a special issue of the TMS publication JOM, March 2022.IntroductionAdditive manufacturing (AM), a family of layer-by-layer digital manufacturing processes, delivers ample opportunities for fabricating near-net-shape or final shape parts with complex internal and external geometries, high freedom in the design, manufacture of customized parts on demand, and consequently have drawn much attention from various industrial sectors [1].

Among all the AM methods, the laser-based methods, i.e., laser powder bed fusion (L-PBF), and laser powder directed energy deposition (LP-DED) are the most widely used ones [2]. In the L-PBF process, there is a powder bed in which a re-coater provides a new layer of powder with each pass and the powder particles are melted locally in a defined pattern by a (few) high-energy laser source(s) and fused to the previously fused layers. In the LP-DED process, the powder is blown to the melt pool, and a new layer of material is fabricated. The L-PBF process is typically limited to a single-powder feedstock; however, provides more freedom in the design of near-net-shaped parts with higher geometrical resolution than the LP-DED process [3].

Conversely, LP-DED is suitable for multi-powder feedstock, functionally graded components, and the fabrication of large parts since there is no restriction of the build box [4].902It has been shown that the AM materials regardless of the manufacturing technique (e.g., L-PBF or LP-DED) exhibit comparable, and in some cases superior/inferior, static mechanical properties as compared to the conventionally manufactured (CM) counterparts [5,6]. This is ascribed to the rapid cooling/solidification rate of AM processes leading to a finer microstructure in AM materials as compared to that of the CM ones, such as casting. Nevertheless, the AM materials experience high melting and cooling/solidification rates, which results in the formation of defects such as pores, lack of fusion (LoF) defects, and inclusions, high surface roughness, and residual stress, all of which affect their fatigue resistance [7,8].

In addition, it has been reported that there is a significant variation in the cooling/solidification rate of L-PBF and LP-DED methods; L-PBF results in three orders of magnitude higher cooling rate than LP-DED [3]. This causes variation in the thermal history a part experiences during fabrication via each of these AM processes, which results in different micro-/defect- structure, and consequently, differences in the mechanical properties.Babuska et al. [3] has recently compared the microstructure and mechanical properties of CoCr alloy manufactured via L-PBF and LP-DED processes. It has been reported that the L-PBF CoCr specimens have significantly higher tensile strength and ductility than the LP-DED counterparts. It has been shown that the considerably higher cooling rate of the L-PBF process resulted in a much finer microstructure while compared with those of the LP-DED samples.

It has been shown that the considerably higher cooling rate of the L-PBF process resulted in a much finer microstructure while compared with those of the LP-DED samples. Several studies have reported various constituent phases in IN718 fabricated via the two AM techniques due to variation in thermal history. The microstructure of the L-PBF IN718 in non-heat treated (NHT) condition is found to be spheroidal/ellipsoidal γ̋ in γ matrix [9], whereas LP-DED counterparts’ is Laves phase in interdendritic γ matrix [10,11]. Likewise, a lamellar α+ß structure has been reported for the L-PBF Ti-6Al-4V [12], while the LP-DED counterparts have fine lamellar Widmanstatten structure along with grain boundary α phase in the coarse columnar prior ß grains [13].

Therefore, it is essential to understand the differences in the microstructure of the part fabricated with various AM techniques such as L-PBF and LP-DED techniques to establish process-structure-property relationships for various material systems.The 17-4 PH stainless steel (SS) is among the materials adopted for the AM processes and has high corrosion resistance as well as the various combination of strength and ductility depending on the applied heat treatment. However, its microstructure constituents in the NHT condition is still debated. Although Murr et al. [14] reported a martensitic microstructure for the NHT L-PBF 17-4 PH SS, Alnajjar et al. [15] showed evidence of austenite by-passing during solidification of L-PBF 17-4 PH SS due to the high cooling/solidification rate, which results in a ferritic microstructure. It is worth noting that the chemical composition of the 17-4 PH SS powder may have influences on the solidification order and stable phases [16,17].

It is worth noting that the chemical composition of the 17-4 PH SS powder may have influences on the solidification order and stable phases [16,17]. Since the cooling/solidification rate varies between the L-PBF and LP-DED process, the as-fabricated microstructure of L-PBF 17-4 PH SS might be different than that of the LP-DED counterpart.This study aims to investigate and compare the microstructure of 17-4 PH SS manufactured by L-PBF and LP-DED methods. The grain structure and phase constituents in the NHT condition are compared for the L-PBF and LP-DED 17-4 PH SS samples. Further, CALPHAD based phase diagrams are generated to guide the design of heat treatments. The microstructure and texture evolution (i.e., crystallographic orientation of the grains) of L-PBF and LP-DED 17-4 PH SS samples are also compared after different heat treatment steps.

The microstructure and texture evolution (i.e., crystallographic orientation of the grains) of L-PBF and LP-DED 17-4 PH SS samples are also compared after different heat treatment steps. The results obtained are fundamental to establish the process-structure relationships for the AM 17-4 PH SS.903Experimental ProceduresMaterial and FabricationThe 17-4 PH SS samples in this study were fabricated via both L-PBF and LP-DED processes, using Ar-atomized 17-4 PH SS powder. The detailed chemical composition of the 17- 4 PH SS powder used in each of the AM processes is listed in Table 1. Table 1. Chemical composition for 17-4 PH SS powder used for fabrication.CCrNiCu MnSiNb MoNOPSNb+TaL-PBF(Wt. %)0.0115.804.603.670.510.320.320.210.020.040.0350.0140.27LP- DED(Wt.

%)0.0115.804.603.670.510.320.320.210.020.040.0350.0140.27LP- DED(Wt. %)0.0116.394.173.320.060.780.270.040.010.020.000.000.27The L-PBF 17-4 PH SS samples were fabricated using an EOS M290 machine with the EOS default process parameters adopted from Ref. [18] under nitrogen shielding gas. The LP- DED samples were fabricated as round bars on a RPM Innovations (RPMI) 557 machine. The parameters were as follows: power of 1070 W, layer height of 0.38 mm, travel speed of 1,016 mm/min and powder feed rate of 15.1 grams/min. The environment was fully inert using Argon for the LP-DED samples.Phase DiagramsTo guide the choice of an optimal stress relief (SR) temperature, Thermo-Calc. software was used to generate the Ni-Cr binary phase diagram for17-4 PH SS. For this purpose, the TCFE9 thermodynamic database for various types of Fe-based alloys and steels such as stainless steels was employed [19,20].

For this purpose, the TCFE9 thermodynamic database for various types of Fe-based alloys and steels such as stainless steels was employed [19,20]. Besides, to incorporate the influence of all the alloy elements in generating the phase diagram, the Nieq (Nickle equivalent) and Creq (chromium equivalent) were obtained following the Schaeffler equations [21]:(wt%)=%Ni+0.5(%Mn)+0.3(%Cu)+25(%N)+30(%C)Nieq Creq(wt%)=%Cr+2(%Si)+1.5(%Mo)+1.75(%Nb)Eq.1 Eq.2The Nieq and Creq values were further calculated for the L-PBF to be 7.6 and 16.6wt%, respectively. For the LP-DED 17-4 PH SS the Nieq and Creq values were found to be 6.1 and 17.7wt%, respectively. Figure 1 shows the binary phase diagram generated by Thermo-Calc. software for the L-PBF and LP-DED samples encountering their corresponding Nieq and Creq values. The Cr amount for the L-PBF and LP-DED samples are indicated on the phase diagrams to predict the phase constituents in various temperatures.

The Cr amount for the L-PBF and LP-DED samples are indicated on the phase diagrams to predict the phase constituents in various temperatures. According to the phase diagram, the SR heat treatment was carried out at 650 °C for 1 hour for both L-PBF and LP-DED samples. A similar SR procedure has been performed on AM 17-4 PH SS in the literature [22,23]. However, it must be noted that conducting heat treatment at this temperature for longer durations may cause austenite reversion on the lath martensite boundaries, which will remain in the microstructure upon cooling [24].904FeBal.Bal.Figure 1. The Ni-Cr binary phase diagram samples generated by Thermo-Calc. software using TCFE9 thermodynamic database [20].Heat Treatment and Microstructure CharacterizationThe heat treatment cycles applied on the samples are shown in Figure 2 schematically.

software using TCFE9 thermodynamic database [20].Heat Treatment and Microstructure CharacterizationThe heat treatment cycles applied on the samples are shown in Figure 2 schematically. Two sets of heat treatments were conducted; in one set, samples were heat treated directly after fabrication, and for the other set, SR at 650 °C for 1 hour followed by air cooling was carried out prior to conducting the heat treatment. Disk-like samples shown in Figure 2(a) were cut in half for a side-by-side comparison between the heat treatment conditions with/without SR. Among the standard heat treatments recommended for wrought 17-4 PH SS, the CA-H1025 heat treatment cycle has been shown to result in both good strength and reasonable ductility as well as high fatigue resistance, and is therefore used in this study.

However, the choice of heat treatment is varied upon on the properties needed for the end use application.The CA (i.e., condition A) annealing step at 1050 °C (1922 F) is shown in phase diagrams (see Figure 1) for both the L-PBF and LP-DED 17-4 PH SS to be in the full austenitic region. Samples were air-cooled after the CA step to room temperature and further aged at 552 °C (1025 F) for 4 hours, i.e., CA-H1025. There are debates and efforts made to show the effect of direct aging in AM martensitic SSs such as Maraging steels; however, it is found that solution annealing heat treatment enhances the mechanical properties significantly, which is attributed to a more homogenized microstructure after CA. It has been reported that direct aging at 552 °C for the L-905PBF 17-4 PH SS fabricated under argon shielding gas results in an abnormal fatigue crack growth (FCG) behavior due to the presence of columnar ferrite grains and string δ-ferrite on the grain boundaries [25].Figure 2.

The schematics of (a) the samples for heat treatment and microstructure characterization, and (b)-(d) heat treatment procedures investigated in this study.The samples were ground and polished using sandpapers with grits 320-4000 following by a mirror-finish polishing step using Chemo-met along with 0.2 µm colloidal silica suspension. The microstructure was further characterized on the plane parallel (X-Z) to the build direction as shown in Figure 2 using a Zeiss 550 crossbeam scanning electron microscope (SEM) with an electron backscatter diffraction (EBSD) detector. The samples were etched using Braha’s reagent to observe the microstructure via SEM. However, for the EBSD analysis as well as electron channeling contrast imaging (ECCI), samples were mirror-finished using a vibratory polisher for 2 hours. The microhardness testing was performed on all the specimens using Anton Par instrument.

The microhardness testing was performed on all the specimens using Anton Par instrument. The force and dwell time for the hardness testing were 100kgf and 5 seconds, respectively.Results and DiscussionNon-Heat Treated (NHT) MicrostructureThe 17-4 PH SS has been reported to have a ferrite-based microstructure while manufactured via AM process in contrast to the precipitated martensitic microstructure in the wrought form. This can be ascribed to the austenite by-passing during fabrication due to the high cooling rate of AM processes [15]. Although the AM processes induce a very high cooling rate naturally, the cooling rate in different AM methods (e.g., L-PBF vs. LP-DED) is not similar. Therefore, the thermal history experienced by the part may be different and result in variation in the microstructure.906The etched microstructure and the ECCI images of L-PBF and LP-DED 17-4 PH SS samples in NHT condition are presented in Figures 3(a)&(b) and (c)&(d), respectively.

The microstructure of L-PBF 17-4 PH SS is significantly different than that of the LP-DED counterpart. It can be seen in Figures 3(a) and (b) that both fine equiaxed and columnar ferrite grains with the lath martensite are present in the L-PBF 17-4 PH SS sample. Normally, the coarse columnar grain structure is seen for most of the AM materials. However, it has been reported that fabricating the L-PBF 17-4 PH SS under nitrogen shielding gas may refine the grain structure and form equiaxed grains as compared to when argon is used as the shielding gas [18]. Therefore, the fine grain structure of the L-PBF 17-4 PH SS sample seen here may be due to the fabrication under nitrogen shielding gas. On the other hand, as presented in Figures 3(c) and (d), a coarse columnar ferritic matrix with Widmanstätten ferrite grains is seen in the microstructure of LP-DED 17-4 PH SS sample in NHT condition.

The coarse grain structure of the LP-DED 17-4 PH SS sample as compared to that of the L-PBF counterpart is ascribed to the significantly lower cooling rate in the LP-DED process than the L-PBF [3]. In addition, the Creq/Nieq ratio is higher for the LP-DED 17- 4 PH SS (~2.9) as compared to that of the L-PBF counterpart (~2.1), which results in higher stability of ferrite in LP-DED sample than L-PBF one.Figure 3. SEM and ECCI micrographs of (a)&(b) L-PBF, and (c)&(d) LP-DED 17-4 PH SS samples.Heat Treated (HT) Microstructure: Solution annealed + agedThe inverse pole figure (IPF) maps of the L-PBF and LP-DED 17-4 PH SS samples in NHT condition are presented in Figures 4(a) and (c), respectively. The EBSD results are taken from the XZ-plane parallel to the build direction as it is schematically illustrated. Note that the colors on the IPF maps represent the crystallographic orientation of the grains.

Note that the colors on the IPF maps represent the crystallographic orientation of the grains. It is seen that the907L-PBF 17-4 PH SS sample (Figure 4(a)) has a significantly finer grain structure than that of the LP-DED counterpart (Figure 4(c)) in NHT condition, consistent with Figure 3. There are grains with size ~1 mm observed in the LP-DED 17-4 PH SS microstructure, although finer grains are also present. These fine grains may have formed surrounding the melt pool boundaries [26]. Considering both the coarse and fine grains, the average grain size for the NHT LP-DED samples is calculated ~93.1 µm. On the other hand, the average grain size for the NHT L-PBF samples is calculated to be ~6 µm.The microstructures of L-PBF and LP-DED 17-4 PH SS samples after SR heat treatment are shown in Figures 4(b) and (d), respectively. As expected, the SR heat treatment conducted on L-PBF and LP-DED 17-4 PH SS samples does not change the grain structure when compared to the NHT condition.

As expected, the SR heat treatment conducted on L-PBF and LP-DED 17-4 PH SS samples does not change the grain structure when compared to the NHT condition. The average grain size after SR is calculated ~101.6 µm for the LP-DED samples, and ~5.7 µm for the L-PBF ones, where the variations from the grain size in NHT condition are minimal. In addition, the SR heat treatment applied herein does not change the hardness of the material; the hardness for the NHT L-PBF sample is measured 327±8 HV, while is measured 319±12 HV for the SR L-PBF counterpart. Similarly, the hardness for the NHT LP- DED sample is measured 353±21 HV, while is measured 334±11 HV for the SR LP-DED counterpart.Figure 4. IPF maps of 17-4 PH SS samples in NHT and SR conditions: (a)&(b) L-PBF, and (c)&(d) LP-DED.It can be seen in Figure 5(a) that the ferritic microstructure of L-PBF 17-4 PH SS observed in Figure 4(a) transforms to martensitic microstructure after conducting CA-H1025 heat treatment.

The phase constituent at 1050 °C (i.e. CA step) is fully austenite (see Figure 1), which transforms to martensite after cooling to room temperature as the martensite start temperature (i.e.,908Ms) for the 17-4 PH SS higher than the room temperature (Ms=150 °C) as reported in Ref. [27]. The hardness is increased (415±9 HV) as compared to the NHT condition (327±8 HV) which can be ascribed to the martensitic microstructure and formation of Cu-enriched precipitates. It is worth mentioning that in case there were any δ-ferrite formed during fabrication, it will remain in the microstructure and has been reported to be along with the lath martensite. However, it is postulated that due to using nitrogen as the shielding gas during manufacturing of L-PBF samples, δ-ferrite might not have formed, or at least formed in a very low fraction.

However, it is postulated that due to using nitrogen as the shielding gas during manufacturing of L-PBF samples, δ-ferrite might not have formed, or at least formed in a very low fraction. It has been well established in the laser welding community that nitrogen decreases the δ-ferrite content in the microstructure of various grades of stainless steels (e.g., 316LN) [28,29]. Conducting SR prior to CA-H1025 (see Figure 5(b)) is found to have minimal influence on the phase constituent and grain structure of L- PBF 17-4 PH SS when compared with the sample undergone only CA-H1025 heat treatment condition (Figure 5(a)). However, it seems that the SR-CA-H1025 L-PBF sample possesses a slightly finer microstructure than that of the CA-H1025, and the hardness is slightly increased (440±7 HV).

However, it seems that the SR-CA-H1025 L-PBF sample possesses a slightly finer microstructure than that of the CA-H1025, and the hardness is slightly increased (440±7 HV). The SR heat treatment may have induced recrystallization; therefore, following heat treatment at CA-H1025 after SR resulted in newly formed finer austenite grains, which then formed finer lath martensite as compared to that of CA-H1025 sample.The microstructure of LP-DED 17-4 PH SS in CA-H1025 (419±14 HV) and SR-CA- H1025 (410±14 HV) heat treatment conditions are presented in Figures 5(c) and (d). As seen, the grain structure is more or less similar to that of NHT and SR conditions (see Figures 3(b) and (d)) although the hardness increased due to the formation of Cu-enriched precipitates.

As seen, the grain structure is more or less similar to that of NHT and SR conditions (see Figures 3(b) and (d)) although the hardness increased due to the formation of Cu-enriched precipitates. The fraction of low angle boundaries increased significantly after CA-H1025 heat treatment compared to the NHT condition indicating the formation of subgrains in the larger grains as a result of dislocation rearrangement in the microstructure. This shows that the CA step’s temperature and duration (i.e., 1050 °C/0.5 hr) is not sufficient for the LP-DED 17-4 PH SS to recrystallize the grains and result in grain refinement. Comparing the microstructure of L-PBF 17-4 PH SS with that of LP-DED one in CA-H1025 condition, the L-PBF sample has fine lath martensite in fine prior austenite grains (shown by black lines), whereas the ferritic microstructure of the LP-DED sample largely remained.

This may be due to the higher Creq/Nieq ratio of LP-DED 17-4 PH SS (~2.9) as well as lower cooling rate in fabrication as compared to the L-PBF process (Creq/Nieq= 2.1), which increases the stability of the ferrite phase in the microstructure of LP-DED 17-4 PH SS samples.909Figure 5. EBSD micrographs of 17-4 PH SS samples undergone CA-H1025 and SR-CA-H1025 heat treatment conditions: (a)&(b) L-PBF, and (c)&(d) LP-DED.Texture EvolutionThe solidification texture for the AM materials is usually reported to be <001> fiber texture, which is the easy growth direction in cubic lattices (i.e., body-centered cubic (BCC), face- centered cubic (FCC)). However, in some cases, the variation in thermal history caused by changing the process parameters (e.g., laser power, scan speed, or build platform preheating, etc.) may alter the crystallographic orientation of the grains [30].

Moreover, the post heat treatment at high temperatures also changes the as-solidified texture due to the formation of new grains/phases [31]. To understand and compare the crystallographic texture of L-PBF and LP-DED 17-4 PH SS samples, their corresponding orientation distribution functions (ODFs) maps for each condition are presented in Figure 6. The ODFs represent the frequency distribution of the continuum of orientations in Euler space. The ODFs intensities are indicated as multiple random density (mrd) units. It is worth noting that only a subspace of Euler space (0≤φ1, Φ, φ2≤90) is required to represent the entire texture information of cubic materials.

It is worth noting that only a subspace of Euler space (0≤φ1, Φ, φ2≤90) is required to represent the entire texture information of cubic materials. The φ2=45° slice of this subspace contained important textures for BCC materials (see Figure 6(a)) and is used to analyze the 17-4 PH SS samples in this study.It can be seen for the NHT L-PBF 17-4 PH SS sample (see Figure 6(b)) that there is a strong cube texture with (001)[010] component; however, there are also components of α-fiber and γ-fiber texture present. The SR heat treatment slightly decreases the texture intensity and randomizes the texture; however, slightly intensified the α-fiber texture with (110)[110] components. Comparing the texture of the CA-H1025 sample with that of the NHT condition, the texture components of CA-H1025 resemble those of NHT ones.

Comparing the texture of the CA-H1025 sample with that of the NHT condition, the texture components of CA-H1025 resemble those of NHT ones. On the other hand, conducting SR910prior to the CA-H1025 (i.e., SR-CA-H1025) results in a weaker γ-fiber texture than that of the CA- H1025 condition, and the absence of the cube texture component as compared to the CA-H1025 specimen. This may be due to the fact that SR induces the recrystallization of the new grains. The L-PBF 17-4 PH SS shows to have the typical texture of the BCC materials, i.e., γ-fiber (i.e.,{111}<110>) [31]. In the case of LP-DED 17-4 PH SS shown in Figure 6(c), there is not a distinct fiber texture present in the NHT condition, and the texture intensity is also very high. This is due to the presence of very large grains in the microstructure (see Figure 4 and Figure 5); therefore, there are not enough grains to show the distributed texture in ODFs.

This is due to the presence of very large grains in the microstructure (see Figure 4 and Figure 5); therefore, there are not enough grains to show the distributed texture in ODFs. It can be seen that CA-H1025 and SR-CA-H1025 result in γ-fiber texture, and γ-fiber + rotated cube texture (i.e., {001}<011>), respectively. Nonetheless, the γ-fiber texture is apparently the dominant texture component in both L-PBF and LP-DED 17-4 PH SS, which also is the typical texture for the BCC materials.Figure 6. (a) Schematic illustration of the important texture components in BCC materials for φ2 = 45°, ODF plots at φ2 = 45° showing texture for (b) L-PBF 17-4 PH SS, and (c) LP-DED 17-4 PH SS after various heat treatments.911ConclusionsThis study presented a side-by-side comparison of the microstructure of the L-PBF 17-4 PH SS with that of the LP-DED counterpart in the NHT condition and after heat treatments.

SR heat treatment temperature was determined by the chemical compositions of the powders and the Ni-Cr binary phase diagrams. In addition, the texture evolution after heat treatment was also investigated using ODFs. The following conclusions are drawn based on the results presented:1. The microstructure of the L-PBF 17-4 PH SS sample was composed of fine ferrite grains along with lath martensite, whereas very coarse ferrite grains with the Widmanstätten ferrites along the grain boundaries were observed for the LP-DED 17-4 PH SS sample. This variation in the microstructure was attributed to the lower cooling/solidification rate as well as the higher Creq/Nieq ratio of the LP-DED 17-4 PH SS as compared to the L-PBF one.2. Using the phase diagram generated by Thermo-Calc., it was found that heat treatment at 650 °C for an hour may be appropriate for the SR heat treatment schedule for both the L- PBF and LP-DED 17-4 PH SS samples.

The microstructure and hardness of stress relieved sample were not changed considerably as compared to those of NHT ones.3. The L-PBF 17-4 PH SS were found to have a cube, partially α- and γ- fiber textures in NHT condition. However, there was no specific texture component characterized for the LP-DED sample in NHT condition based on the EBSD scans performed which can be ascribed to its very large grains.4. Conducting stress relief (SR) prior to the heat treatment (i.e., SR-CA-H1025) resulted in a less texturized microstructure. This may be due to the effect of SR on inducing recrystallization upon the following heat treatment and result in newly formed grains with random textures.AcknowledgmentThis paper is based upon the work partially funded by the National Aeronautics and Space Administration (NASA) under Award #80MSFC19C0010.

MPARING MECHANICAL AND GEOMETRICAL PROPERTIES OF LATTICE STRUCTURE FABRICATED USING ELECTRON BEAM MELTINGSang-in. Parka, David W. Rosena, and Chad E. Dutyba The G. W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332 b Oak Ridge National Laboratory, Oak Ridge, TN 37831ABSTRACTTo design lattice structure, a uniform voxel based approach is widely used which divides a part into unit volumes (e.g., cubes) and maps lattice topology into those volumes. In contrast, conformal lattice structures represent a second design method for constructing lattices in which unit cells are constructed parallel to the surface to be reinforced and are deformed in a manner that enables them to conform to the surface. In this paper, the strength of lattice structures designed using these two methods (uniform voxel based and conformal) are compared based on additive manufacturing (AM) process effects.

In this paper, the strength of lattice structures designed using these two methods (uniform voxel based and conformal) are compared based on additive manufacturing (AM) process effects. For this purpose, spheres filled with three types of lattice structure are fabricated using electron beam melting technology and tested in compression. Effects of AM processes are studied in two ways – volumetric and structural performance equivalence. Struts in lattice structures are observed through a microscope to examine volume- equivalence and tests are simulated numerically and compared to identify structural equivalence.1INTRODUCTIONCellular materials such as foam, honeycomb, and lattice structure are used in applications due to their special mechanical properties which cannot be achieved by conventional bulk material. The use of cellular material expands the design space for mechanical properties [1].

The use of cellular material expands the design space for mechanical properties [1]. Generally, cellular materials can be tailored for high strength to weight ratio, thermal conductivity and energy absorbance [2]. A lattice structure is one type of cellular material which is comprised of a connected network of struts. Among the cellular material, the lattice structure has a distinguished characteristic, which is lattice structures are composed of representative unit cells that define their geometries and topologies. This enables engineers to design mechanical properties of lattice structures such as elastic modulus, yield strength and fracture strength for specific applications [1, 3].There are two approaches for designing the lattice structures – a uniform and a conformal lattice approaches [4]. In the uniform approach, the volume of a part is divided into small regular blocks and the topology of unit cell is mapped into the blocks.

In the uniform approach, the volume of a part is divided into small regular blocks and the topology of unit cell is mapped into the blocks. Lattice structures can be simply obtained through this process and struts inside lattice structures are fully connected. However, obtained lattice structures often become broken or unconnected near surfaces as Figure 1 (a).(a) Uniform lattice approach(b) Conformal lattice approachFigure 1 Design approach for lattice structureThe conformal lattice structure approach is proposed to solve the problem of broken struts at part boundaries [5]. In the conformal lattice approach, conformal hexahedral meshes on the target surfaces are created and unit cells are mapped into the meshes.

In the conformal lattice approach, conformal hexahedral meshes on the target surfaces are created and unit cells are mapped into the meshes. Since meshes from the mapping process conform to target surfaces, the struts are fully connected as shown in Figure 1 (b)Although lattice structures generally give better properties than other cellular materials, complex geometries in lattice structures make it difficult [6, 7] or impossible to manufacture lattice structures using conventional manufacturing methods. Additive manufacturing (AM) technologies can be a solution for the problems in manufacturing lattice materials. Since AM processes deposit materials layer by a layer based on slicing information of parts, lattice structures can be easily fabricated without a consideration for tool interference. Through the past two decades, various AM processes have been used to fabricate lattice structures.

Through the past two decades, various AM processes have been used to fabricate lattice structures. The Electron Beam Melting (EBM) process is one such AM process which uses an electron beam to melt metal powder and form the shape of part cross-sections. Recently, this process has been applied to build lattice structure with Ti6Al4V metal powder [8, 9].Fabricating lattice structure takes advantage of AM processes. However, the processes introduce shape variations which are geometrical fluctuation or errors of fabricated features compared to their designed shapes and sizes. Since the AM processes deposit material based on cross sectional layer information, fabricated parts have stair steps [10, 11]. Additional factors cause geometric deviations in the EBM process. Since metal powder is selectively melted by the electron beam, the unmelted powder near melt pools can become stuck on the part surfaces. This phenomenon makes the surfaces rough and uneven. Yang et al.

This phenomenon makes the surfaces rough and uneven. Yang et al. reported that struts in fabricated lattice blocks show dimensional variations due to Ti6Al4V powder particles [12]. Parthasarathy et al. evaluated EBM processed lattice blocks using 3D digital reconstructions and showed variations in dimensions of the blocks [13]. Recently, List et al. studied the relationship between process parameters and geometrical and mechanical parameters [14]. Previous researches indicate that AM processes affects geometrical dimensions as well as mechanical properties of manufactured parts.The goal of this paper is to study the effects of the EBM AM process on geometrical and mechanical properties of fabricated lattice structure.

To achieve the goal, two research questions are investigated; which design method between the uniform and conformal lattice approach gives better mechanical properties? and how to quantitatively measure shape variations resulting from the additive manufacturing process? We hypothesize that conformal lattice structure will have higher fracture strength for the first question and geometrically and mechanically equivalent values of strut diameters can be used for measuring shape variations for the second question. In this paper, three tasks are conducted in order to test hypotheses. Related to the first research question, fracture strengths of meshed balls fabricated by EBM process are compared, which are designed by the two design approaches.

Related to the first research question, fracture strengths of meshed balls fabricated by EBM process are compared, which are designed by the two design approaches. Related to the second research question, the mass- equivalent strut diameter is determined for meshed balls based on microscopic observation and design tools and the response-equivalent strut diameter is calculated based on comparison of compression tests and numerical analysis.2 FABRICATION OF SPECIMENS2.1 Design of SpecimensAs specimens, four kinds of meshed balls were designed by uniform and conformal approaches. Schematic procedures for designing meshed balls are shown in Figure 2. The main difference between the two approaches is the meshing process. In the uniform lattice, regular meshes fill the volumetric region of the balls and the meshed balls have unit cells mapped into these regular meshes.

In the uniform lattice, regular meshes fill the volumetric region of the balls and the meshed balls have unit cells mapped into these regular meshes. However, in the conformal lattice, meshes are generated on selected base surfaces and meshed balls are constructed by generating a layer of hexahedra mesh elements, then mapping the unit cell into the elements. For the uniform lattice, the commercial software package, Magics, was used. For the conformal lattice, the TrussCreator software package was used, which is a plug-in for the Siemens NX CAD system [15].(a) Uniform lattice procedureRotationBase surface(b) Conformal lattice procedure Figure 2 Schematic procedure of lattice design procedure(a) Cube(b) Diamond(c) PentahedronFigure 3 Unit cell typeThe meshed balls are comprised of three types of unit cells, which are presented in Figure 3. Designs and specifications of the meshed balls are listed in Table 1.

Designs and specifications of the meshed balls are listed in Table 1. Design 1 is designed by the uniform lattice procedure and Designs 2, 3 and 4 are constructed by the conformal lattice procedure. Design 1 and 2 are composed of the diamond unit cell. Design 3 and design 4 are comprised of cube and pentahedron unit cells, respectively. To keep the mass of all mesh balls similar, the strut diameters in the meshed balls were not kept constant, but were adjusted. The design strut diameters of designs 1 and 2 are 0.4 mm and 0.35 mm, respectively. The design strut diameter of design 3 is 0.4 mm and 0.35 mm is used for design 4.

The design strut diameters of designs 1 and 2 are 0.4 mm and 0.35 mm, respectively. The design strut diameter of design 3 is 0.4 mm and 0.35 mm is used for design 4. The volume is calculated in Magics.Table 1 Design of meshed ballsDesignWhole designCross sectionDesign approach / Unit cell typeStrut Diameter (mm)Volume (mm3)1Uniform / Diamond0.412102Conformal / Diamond0.3512323Conformal / Cube0.410414Conformal / Pentahedron0.351080It is worth noting that the conformal lattice balls have one complete layer of lattice structure, while Design 1, with uniform lattice, has one layer in most regions, but this layer is truncated, as seen in the upper left and lower right regions of the cross section view in Table 1.2.2 Fabricated SpecimensNine meshed balls (two of designs 1, 2, 3 and three of design 4) were built in the Arcam A2 at Oak Ridge National Laboratories using Inconel 718 metal powder.

The “Net” theme was used in the Arcam 3.2 EBM control software for melting the powder. The beam current was 2.5 mA and the beam speed was 300 mm/s. The fabricated meshed balls are shown in Figure 4. The fabricated masses are compared with estimated masses using the reference density (0.00819 g/mm3) of Inconel 718 and reported in Table 2. The masses are at least 145% more than that estimated for the designed balls. The unmelted powder stuck on strut surfaces and manufacturing tolerance generated variations in the geometrical dimensions and increased the size of the struts. These lead to more mass in the meshed balls.

These lead to more mass in the meshed balls. The amount of increases is quantified in the later sections.(b) Design 1(c) Design 2(a) Specimens on the build plate(d) Design 3(e) Design 4Figure 4 Fabricated meshed ballsTable 2 Comparison of massUnit: gDesign 1Design 2Design 3Design 4Estimated mass in design Fabricated mass Relative difference (%)9.91 28.15 +184%10.09 24.7 +145%8.53 21.7 +154%8.84 26.6 +193%3 COMPRESSION TESTTo investigate the mechanical properties of mesh balls, we conducted compression tests. The build direction in the EBM machine was used as the axis for the compression tests. Representative force-displacement curves from the compression tests are shown in Figure 5 and the maximum loads and corresponding displacements are listed in Table 3. For comparison purposes, the maximum loads are normalized by mass of meshed balls.

For comparison purposes, the maximum loads are normalized by mass of meshed balls. Design 4 composed of pentahedral unit cells carried the highest forces that is 147% more than design 1, which was designed by the uniform lattice approach.There are two noticeable points. The first is that design 1 endured higher force than designs 2 and 3. This means that our first hypothesis cannot be validated by this particular example. Although the conformal lattice approach gives more possibilities to design higher strength lattice structure, the strength of a lattice structure is also governed by details of its geometrical connectivity. The conformal lattice approach insures that lattice structures are fully connected in a layer. However, this does not guarantee that the connectivity is always aligned to a direction of the external load.

However, this does not guarantee that the connectivity is always aligned to a direction of the external load. In case of design 4, since many struts are aligned with the compressive force direction, the measured strength is much higher than other meshed balls.Next, the force-displacement curve of design 3 indicates that struts in the meshed ball are buckled. This can be observed in the crashed geometry. Since there are no internally connected struts in the cubic unit cell, the buckling mode limits the strength of cubic unit cell lattice structure although the arrangement of struts are parallel to the external force.

On the other hand, other meshed balls comprised of diamond and pentahedral unit cells show fracture mode failure.0100020003000400050006000700002468101214Force(N)Displacement (mm)Design 3Design 2Design 4Design 1Figure 5 Force-displacement curve during testBuckling(a) Design 1(b) Design 2(c) Design 3(d) Design 4Figure 6 Specimens after compression testTable 3 ultimate force and corresponding displacementDesign1 2 3 4Ultimate force (N) 5034.9 2554.1 3512.4 6844.0Force/Mass (N/g) 178.86 103.40 161.86 264.59Displacement at ultimate force (mm) 7.47 6.53 6.89 8.704 DETERMINING MASS-EQUIVALENT STRUT DIAMETERIn order to investigate shape variations due to the EBM process, the selected fabricated balls of designs 1 and 4 are observed through the microscope. Figure 7 shows representative strut surfaces from two orientations.

Figure 7 shows representative strut surfaces from two orientations. When observed in the direction parallel to the build direction, the surfaces are smooth without unmelted powder and stair-steps can be detected based on boundary of the melting pool. However, in the direction perpendicular to building direction, lots of unmelted powder is stuck on strut surfaces. The reason for this phenomenon is that the electron beam melts powder in the top layer of the powder bed and powder particles below the top layer partially melt or otherwise adhere to the melt pool and protrude from the strut surfaces. As a result, the unmelted powder produces shape variations and rough surfaces.The fabricated strut diameters are measured in digital images taken from a microscope. Table 4 lists minimum, maximum and average value of measured strut diameters.

Table 4 lists minimum, maximum and average value of measured strut diameters. In the case of design 1, the maximum and minimum values are 0.79 mm and 0.452 mm, respectively, while the design strut diameter is 0.4 mm and for design 4 which has a designed diameter of 0.35 mm, the maximum and minimum values are 0.737 mm and 0.377 mm. Based on observation, it is noted that large variations in dimension exist due to the EBM process.

The mass-equivalent strut diameter is the diameter which gives the same mass in the designed ball as in the fabricated part. To find the mass-equivalent strut diameter, parametric studies were conducted. Meshed ball volumes were calculated in Magics. Based on the strut diameter and corresponding volume of meshed ball design, regression models were generated. The mass-equivalent strut diameters are summarized in Table 5. Compared to designed strut diameters, the mass-equivalent strut diameters are about 1.8 times greater. The estimated masses of meshed balls using the mass-equivalent strut diameters show less than 3% relative error with respect to mass of the fabricated balls.

The estimated masses of meshed balls using the mass-equivalent strut diameters show less than 3% relative error with respect to mass of the fabricated balls. This means that the fabricated dimensions are shifted from the design due to shape variation during the EBM process and the deviations seem to be systematic and repeatable.5 DETERMINING RESPONSE-EQUIVALENT STRUT DIAMETERMass-equivalent strut diameters can be used for measuring mismatch in geometrical aspects such as the mass and volume between a desired lattice and a fabricated lattice. However, fabricated lattice structures cannot support the mechanical load that would be expected from a lattice with the mass-equivalent strut diameter since partially melted powder particles are partially bonded to the strut but do not necessarily contribute to its mechanical properties.

Microscopic observations in the previous research show there exist critical defeats in the cross section of struts which degrade the structural performance.[14] In order to quantify the effects of shape variation due to the EBM process on the mechanical response, the response-equivalent strut diameter is introduced and calculated in this section.To simulate compression test in the previous section, the geometrical information of conformal meshed balls are exported to ABAQUS. Figure 8 (a) shows the finite element analysis (FEA) model of design 2 for finite element analysis. One-eighth model of meshed ball is used and symmetric boundary conditions are applied. Shear flexible beam elements are used tomodel struts. To describe plastic deformation during compression tests, the plastic constitutive model of Inconel 718 is applied and shown in Figure 8 (b). For elastic deformation, the elastic modulus is set to 208 GPa and 1210 MPa is assigned for the plastic behavior and yield strength.

For elastic deformation, the elastic modulus is set to 208 GPa and 1210 MPa is assigned for the plastic behavior and yield strength. Frictionless contact is assumed at the interface between the meshed ball and the grip. The grip is moved upward as 4mm so that 8mm of the displacement in total is simulated.To compare the responses of numerical analysis with the test result, the design strut diameters are assigned to each strut in the three meshed balls. Figure 9 describes the force – displacement history of meshed balls. The responses follow the trends of the test but there are large differences in absolute values of forces which the balls support. Table 6 lists ultimate forces during simulations. The ultimate force from FEA is significantly lower than the test results. It is worth noting that the overshoot is observed in case of design 3. This is because in this simulation the buckling phenomenon is not formulated.

It is worth noting that the overshoot is observed in case of design 3. This is because in this simulation the buckling phenomenon is not formulated. Therefore, the peak in the design 3 curve near a displacement value of 6 mm of Figure 9 was selected as the ultimate force. The force ratios in Table 6 are ratio between ultimate forces in each ball and that of design 2. The ratios from the analysis are close to those from the compression test.In order to find the response-equivalent strut diameter a parametric study was performed by re-running the FEA models for different strut diameters. Figure 10 compares four different force-displacement histories of meshed balls with the compression test result for design 4. Based on the parametric study the regression models were constructed to find response-equivalent strut diameters. The response-equivalent strut diameters were calculated using regression models and listed in Table 7.

Table 8 shows that mass-equivalent strut diameters are about 1.8 times thicker and response-equivalent strut diameters are about 1.25 times thicker than design strut diameters. The trends of the results correspond well with the cross section of a fabricated strut shown in Figure 11 (a). The fully melted region is shown in the core of the strut and the unmelted powder is stuck on the melting region. The powder contributes to the mechanical performance but the contribution is limited. However, the powder increases the dimensions so that mass-equivalent strut diameters are much larger than design strut diameters.

However, the powder increases the dimensions so that mass-equivalent strut diameters are much larger than design strut diameters. The diameters are compared schematically in Figure 11 (b).Table 8 Ratios among diametersDesign 1 Design 2 Design 3 Design 4Mass-equivalence to design 1.81 1.86 1.81 1.82Response-equivalence to design1.24 1.26 1.28Response equivalent strut diameterMass equivalent strut diameterDesignstrut diameter(a) Cross section of a fabricated strut(b) Schematic comparison among diametersFigure 11 Comparison among diameters7 CONCLUSIONIn this paper, several meshed balls were designed and fabricated in order to investigate the effects of design methods and the EBM process on the geometrical and mechanical properties of a lattice structure. The first hypothesis we proposed was not supported since improper selection of unit cells in the conformal lattice approach can weaken the strength of the lattice structure.

The first hypothesis we proposed was not supported since improper selection of unit cells in the conformal lattice approach can weaken the strength of the lattice structure. However, since the conformal lattice approach provides more design possibility such as lattice directions, more durable lattice structures can be designed by the approach, we believe.Related to the second hypothesis, two equivalent-strut diameters can represent the amount of deviation from designed diameters in geometrical and mechanical aspects. The mass-equivalent strut diameter can be used for estimation of volume and the size of the struts after fabrication. The response equivalent strut diameters can be utilized in the evaluation process for changes in mechanical behavior of lattice structure due to the EBM process. Future work is to set the functional relationship between equivalent strut diameters and EBM process parameters.

Future work is to set the functional relationship between equivalent strut diameters and EBM process parameters. The relationship can provide mapping for better fabrication using EBM process.8 ACKNOWLEDGMENTFinancial support for this work has been funded by National Science Foundation (NSF) grant CMMI-1200788. The authors would like to appreciate Larry Lowe and William Sames at Oak Ridge National Laboratory for fabricating specimens. The authors also would like to thank professor Seung-kyum Choi at Georgia Tech for advising.9 REFERENCE 1.Ashby, M., Hybrid Materials to Expand the Boundaries of Material‐Property Space. Journal of the American Ceramic Society, 2011. 94(s1): p. s3-s14. Evans, A.G., J.W. Hutchinson, N.A. Fleck, M.F. Ashby, and H.N.G. Wadley, The topological design of multifunctional cellular metals. Progress in Materials Science, 2001. 46(3–4): p. 309-327. Deshpande, V.S., N.A. Fleck, and M.F. Ashby, Effective properties of the octet-truss lattice material.

MPARISON OF ALSI10MG AND AL 6061 PROCESSED THROUGH DMLSBenjamin A. Fulcher*, David K. Leigh*, Trevor J. Watt†Harvest Technologies, Belton, TX 76513 †Department of Mechanical Engineering, The University of Texas at Austin, Austin, TX 78712AbstractDirect Metal Laser Sintering (DMLS) processing of aluminum alloys has been primarily limited to a casting grade of aluminum, AlSi10Mg. The reasons for the choice of AlSi10Mg by machine manufacturers are presently unknown; however, it is suspected that the reduced coefficient of thermal expansion (CTE) due to the presence of Silicon may enhance DMLS processability. Aluminum 6061 (Al 6061) is a commonly used alloy across a wide range of industries and applications, and Harvest has observed a high interest in DMLS-manufactured Al 6061 products. However, the higher CTE value potentially presents greater challenges in controlling the shrinkage-induced warp common during DMLS.

However, the higher CTE value potentially presents greater challenges in controlling the shrinkage-induced warp common during DMLS. The work presented in this paper was performed in an effort to understand differences in manufacturability as well as mechanical properties of DMLS-processed AlSi10Mg and Al 6061.1. IntroductionCasting, welding, and DMLS of aluminum alloys pose several challenges: First, aluminum rapidly forms an oxide layer when exposed to air which requires strong fluxes or an inert atmosphere. Second, water vapor or any adsorbed hydrogen-bearing species can result in hydrogen porosity during solidification, as molten aluminum has a high solubility limit for hydrogen which diminishes rapidly upon solidification. Third, aluminum has a high thermal diffusivity which requires a high energy input to overcome conductive cooling into the surrounding material.

Third, aluminum has a high thermal diffusivity which requires a high energy input to overcome conductive cooling into the surrounding material. Some aluminum alloys (6xxx series in particular) are also susceptible to hot tearing and hot cracking in casting and welding processes. Fracture can occur in the solid- state due to thermal contraction (hot cracking), or above the solidus temperature due to solidification shrinkage (hot tearing).AlSi10Mg is currently the most commonly used aluminum alloy in DMLS, presumably due to the ease of processing. Aluminum-silicon alloys are primarily used in castings due to their high melt fluidity and relatively low shrinkage [1, 2]. Adding more than a few percent of silicon to aluminum significantly improves these properties, reducing the likelihood of hot tearing during solidification. Lower shrinkage results in reduced strain levels and therefore less susceptibility to cracking during solidification.

Lower shrinkage results in reduced strain levels and therefore less susceptibility to cracking during solidification. Furthermore, if the melt is fluid enough it is able to “heal”, or back-fill, any cracks that form while the material is in its mushy state. Silicon, with a high heat of fusion, contributes to an alloy’s fluid life by releasing heat into the melt during solidification, thereby slowing the solidification and reducing strain rate in the mushy zone.Solidification range (∆T, or Tliquidus - Tsolidus) is also relevant to hot tearing. A larger ∆T results in a greater chance for hot tearing because less liquid is available for interdentritic feeding when the material reaches the solidus or eutectic temperature. Near-eutectic alloys such as AlSi10Mg have a relatively low ∆T (40 K), whereas Al 6061 has a relatively large ∆T (70 K) which is further increased by non-equilibrium solidification and fast cooling rates and in DMLS.

In casting processes, the hot tearing of aluminum alloys is reduced by increasing the mold temperature, and reducing the pour temperature [3]. The solidification range for an AlSi10 alloy is represented in the Al-Si binary phase in Figure 1, and the non-equilibrium solidification behavior of an Al-Mg alloy is displayed in Figure 2.∆TAlSi10Figure 1. Al-Si binary phase diagram illustrating thermal behavior of AlSi10 alloys [4]Figure 2. Non-equilibrium solidification of an Al-Mg alloy with composition C0. Solidification begins with the rejection of α solid-solution crystals. If time is insufficient for diffusion of Mg into the solidified core, microsegregation occurs and ∆T increases [5].Given the potential difficulties associated with processing Al 6061 in DMLS it is not surprising that it is not widely used in the additive manufacturing industry.

However, Al 6061 is commonly used across multiple industries, is often used for CNC prototyping, and Harvest has observed a desire from aerospace customers for DMLS Al 6061 components.2. Results and DiscussionScan track testing was performed for Al 6061, AlSi10Mg microstructure and mechanical test samples were built according to the default EOS parameter set, and Al 6061 samples were built according to the parameters listed in Table 1. Note that two builds were performed in Al 6061, referred to as version 1 (v1) and version 2 (v2). The second build employed a preheat scan in an effort to alleviate cracking issues observed in the first build.Table 1. DMLS processing parameters used in this studyLaser Power (W) Scan Speed (mm/s) Scan Spacing (mm) Layer Thickness (µm) Platform Temperature (ºC)AlSi10Mg EOS_DirectPart EOS_DirectPart EOS_DirectPart 30 30Al 6061 (v1) Al 6061 (v2)370 940 0.19 30 3070 / 370 3,000 / 940 .19 40 802.1.

Al 6061 Scan Track TestingPrior to building test samples in Al 6061, a scan track design of experiments was performed to provide a general idea of appropriate scan speed and laser power. In the first scan track test, laser power (P) was varied from 70 to 370 W and scan speed (S) was varied from 1000 to 3000 mm/s. In the second scan track test, the laser power was varied from 270 to 370 W and the scan speed was varied from 500 to 1500 mm/s. Scan spacing was left at the default EOS value for AlSi10Mg of 0.19 mm. The tracks were examined under microscope and characterized by appearance.

Scan spacing was left at the default EOS value for AlSi10Mg of 0.19 mm. The tracks were examined under microscope and characterized by appearance. The following descriptions explain the characters used in Table 2 and Table 3.U – Un-melted  X – Balling, broken tracks, unconnected tracks  Y – Balling, unconnected tracks  Z – Balling with connected tracks  R – Rough with connected tracks  A – Smooth with connected tracksIn Table 2 and Table 3, the lower-left corner corresponds with the highest laser energy input and the upper-right corner corresponds with the lowest energy input. The best results in the two studies tend toward higher laser energy.Table 2. First Al 6061 scan track testTest #1370295Power (W)220145703000 ) s / 2500 m m 2000 1500 1000( d e e p SX Y Z Z AX Y Y Z RX X X Y YX X X X YU U U U XTable 3.

First Al 6061 scan track testTest #1370295Power (W)220145703000 ) s / 2500 m m 2000 1500 1000( d e e p SX Y Z Z AX Y Y Z RX X X Y YX X X X YU U U U XTable 3. Second Al 6061 scan track testTest #2370345Power (W) 3202952701500 ) s / 1250 m m 1000 750 500( d e e p SZ Z A A AZ Z R A AY R R A AY R R A AY Y R R AAfter performing the scan track testing, laser power of 370 W and scan speed of 940 mm/s were chosen for the first Al 6061 build. This combination ensures good melt characteristics while the relatively high scan speed facilitates a fast build rate. As previously mentioned, a second build was run after the first in order to alleviate cracking issues observed in the first build (see section 2.2). The second build employed a preheat scan of 70 W at 3,000 mm/s which corresponds with an un-melted track in the scan track testing. Figure 3 displays three representative scan tracks from the first test.Figure 3.

Figure 3 displays three representative scan tracks from the first test.Figure 3. (upper) Al 6061 scan track testing at P=220 W, S=1000 mm/s (type Y); (middle) P=370 W, S=2000 mm/s (type Z); (lower) P=370 W, S=1000 mm/s (type A)2.2. Microstructural ComparisonSpecimens were sectioned with a low-speed diamond saw and encapsulated in epoxy. Each epoxy-mounted specimen was sequentially ground using a rotary polisher with 400, 600, 1200 and 4000 grit SiC papers immersed in water, then sequentially polished with 9-μm, 3-μm, and 1-μm diamond suspensions. Final polishing used a 0.04-µm colloidal silica suspension. The specimens were rinsed with water and ethanol and air-dried between each grinding/polishing step.Specimens were etched using one of two techniques: The first technique used 2% aqueous fluoroboric acid (HBF4, “Barker’s reagent”) for approximately 60 seconds, which resulted in tint etching of the AlSi10Mg and precipitate etching of the Al 6061.

The second technique used the same reagent with an electrical current applied at 25 VDC with the specimen as the anode. This electrochemical etching results in grain orientation coloring when viewed with an optical microscope equipped with cross-polarized filters and a λ tint plate. After etching, specimenswere analyzed using a Zeiss Axiovert 405M optical polarizing microscope, and a JEOL JSM- 5610 scanning electron microscope equipped with energy dispersive spectroscopy (EDS).2.2.1. As-built Cross-Section ImagesThe following images were taken of material in the as-built (melted, stress relieved)condition.AlSi10Mg as-built micrographsOptical microscope images of electrolytic etched AlSi10Mg in the as-built condition are shown in Figure 4. In general there is little difference between the different melt pools. Melt pool edges are tinted yellow or blue, with the central region colored magenta.

In general there is little difference between the different melt pools. Melt pool edges are tinted yellow or blue, with the central region colored magenta. The melt pool boundaries are extremely well defined and there is sporadic evidence of epitaxial solidification. In Figure 5 epitaxial solidification is evident in the middle of the melt pools as propagation of the same color, though the melt pool edges tend to show greater orientation contrast.Figure 4. Optical micrographs of as-built AlSi10Mg,electrolytic etch; XY plane (left), and XZ plane (right)Figure 5. Optical micrographs of as-built AlSi10Mg, electrolytic etch; XZ planeSEM images of the XY plane for AlSi10Mg in Figure 6 show a cellular solidification structure with cells approximately 0.4-0.5 μm in size. The cells increase in size up to about 1 μm at the melt pool boundaries.

The cells increase in size up to about 1 μm at the melt pool boundaries. These larger cells are likely HAZ material that has been heated up to just below the melting temperature, resulting in coarsening of the microstructure. This region is surrounded on one side by a 2-3 μm band of detached dendrites/cells. Beyond this band the melt pools develop a “honeycomb” style inter-connected microstructure.Figure 6. SEM images of AlSi10Mg as-built, chemical etch; XY planeIn Figure 7, the cells in the XZ plane show some evidence of epitaxial solidification near the melt pool bottoms, though the sides of the melt pools tend to have non-epitaxial cells aligned in the direction of greatest heat transfer. The detached cells seen in the XY plane are observed on the underside of the coarsened cells. Since these are evident as spots in both the XY and XZ plane, it is apparent that they are actually nodular in shape.

Since these are evident as spots in both the XY and XZ plane, it is apparent that they are actually nodular in shape. Additional analysis is needed to determine if this is due to solidification of the original melt pool or heating from the overlapping melt pool, though it is likely the latter.Figure 7. SEM images of AlSi10Mg as-built, chemical etch; XZ planeAl 6061 (v1) as-built micrographsImages of as-built, polished Al 6061 (v1) samples are shown in Figure 8. Significant cracking and porosity is observed throughout the material. The cracks in the XY plane are arranged in a “mud crack” fashion suggesting biaxial tensile stresses, and are often coincident with pores.Figure 8. As-built, polished Al 6061 (v1) XY plane (left) and XZ plane (right)Figure 9 is an SEM image of a crack in the as-built Al 6061 in the XY plane.

As-built, polished Al 6061 (v1) XY plane (left) and XZ plane (right)Figure 9 is an SEM image of a crack in the as-built Al 6061 in the XY plane. Tendrils that connect opposing crack faces suggest that fracture occurred when this material was still in the “mushy” state, or between the solidus and liquidus temperatures.Figure 9. SEM image of Al 6061(v1) crack in the XY planeIn Figure 10, melt pool boundaries are revealed as well as grain orientation contrast for the as-built Al 6061. Epitaxial solidification is clearly evident for the Al 6061, unlike the AlSi10Mg. Grain orientations span many build layers and can consume large portions of a particular melt pool’s width. Cracks are predominately intergranular in nature – this behavior is expected from the previous figure which showed evidence of hot tearing in the fracture surfaces.

Cracks are predominately intergranular in nature – this behavior is expected from the previous figure which showed evidence of hot tearing in the fracture surfaces. However, unlike traditional hot tearing which is due to macroscopic geometric constraints, the cracks observed here occur on a much smaller length-scale and occur throughout the melt pool. Cracks also tend to correspond with greater changes in grain color (grain orientation), which suggests that fracture is due to anisotropic contraction during solidification and cooling. The cracking is remarkably consistent along the Z axis and spans across multiple powder layers.Figure 10. Optical micrograph of Al 6061 (v1) as built, electrolytic etch; XZ planeAl 6061 (v2) as-built micrographsThe second build of Al 6061 (with preheat scanning) shows marked improvement over the first build, although there is still work to be done to completely eliminate cracking and porosity.

The preheat scan effectively reduces thermal gradients during cooling and thereby reduces thermal cracking. While there are some regions of significant cracking in the second build of Al 6061, most of the material exhibits only fine-scale cracks that are difficult to see without using dark-field imaging.Figure 11. Optical micrographs of Al 6061 (v2) as-built, polished; XZ plane. Bright-field images are shown on the left with corresponding dark-field images on the right.2.2.2. Heat-Treated Cross-Section ImagesThe following images were taken of material in the as-built (melted, stress relieved), andheat treated (HIP, solutionized, aged) condition.AlSi10Mg heat-treated micrographsThe XY plane micrographs of the heat-treated AlSi10Mg in Figure 12 show that the original cellular microstructure has been converted into coarsened nodular particles. Furthermore, the grain orientation no longer follows the melt pools as in the as-built condition.

Furthermore, the grain orientation no longer follows the melt pools as in the as-built condition. The XZ plane micrographs in Figure 13 show some particles aligned along previous cellular growth directions. The grain orientation is irregular, and does not consistently span multiple powder layers as in the Al 6061.Figure 12. Optical micrograph of AlSi10Mg-T6, electrolytic etch; XY planeFigure 13. Optical micrograph of AlSi10Mg-T6, electrolytic etch; XZ planeAl 6061 (v1) heat-treated micrographsFigure 14 shows that grain orientation is largely consistent across the XY plane, and the cracks no longer correspond with significant changes in grain orientation. The micrograph in Figure 15 is similar to the as-built Al 6061 (see Figure 10), though the pool boundaries arelargely absent.

The micrograph in Figure 15 is similar to the as-built Al 6061 (see Figure 10), though the pool boundaries arelargely absent. This confirms that pool boundaries were originally marked by coarsened precipitates (presumably Mg2Si) that have been dissolved and redistributed as smaller dispersed precipitates due to the T6 heat treatment.Figure 14. Optical micrographs of Al 6061-T6, electrolytic etch; XY planeFigure 15. Optical micrograph of Al 6061-T6, electrolytic etch; XZ plane2.3. Mechanical Properties ComparisonTensile bars and Charpy impact samples were built in AlSi10Mg and Al 6061. Tensile bars were built according to ASTM E8 (half size) along the X axis and the Z axis. Hand finishing was performed on the X tensile bars in order to remove support structure residue, and no finishing was performed on the Z tensile bars. Charpy specimens were built and the notch was subsequently machined.

Charpy specimens were built and the notch was subsequently machined. As seen in Table 4, mechanical properties for the Al 6061 specimens arehighly anisotropic. The Z specimens exhibit much higher UTS and impact energies than the X specimens, however even the Z properties are much lower than expected. These results are preliminary and should improve with future parameter optimization efforts. Fracture surfaces of the Charpy samples can be seen in Figure 16. The anisotropy in the Al 6061 is evident in these images as well, and there is little observable difference between the X and Z fracture surfaces for the AlSi10Mg. At the time this report was written, mechanical testing of Al 6061 (v2) had not yet been performed.Table 4.

At the time this report was written, mechanical testing of Al 6061 (v2) had not yet been performed.Table 4. UTS and impact energy for AlSi10Mg and Al 6061 specimensUTS (ksi) Impact Energy (J) ConditionAlSi10Mg, X-Up 48.7 4 HIP, T6AlSi10Mg, Z-Up 43.9 4 HIP, T6Al 6061, X-Up (v1) 6.1 1.5 HIP, T6Al 6061, Z-Up (v1) 33.4 7 HIP, T6Al 6061 (v2) In process In process -Figure 16. Fracture surfaces of AlSi10Mg Charpy SpecimensSEM images of the fracture surfaces of AlSi10Mg-T6 Charpy bars built along the X and Z axes are shown in Figure 17. These surfaces reveal traditional ductile fracture through micro-void coalescence. Once again, there is little observable difference between the X and Z fracture surfaces.Figure 17. SEM images of AlSi10Mg-T6 Charpy fracture surfaces; (left) X-Up and (right) Z-UpThe fracture behavior for the Al 6061 (v1) material is markedly different than that of theAlSi10Mg. This material fails in a manner similar to directional fibrous composites.

This material fails in a manner similar to directional fibrous composites. Sudden changes in the fracture surface orientation and texture seen in Figure 18 are likely due to residual melt pool boundaries, where certain grain orientations change the direction of growth as shown in the previous color metallography images. The columnar fracture surfaces indicate that the original microstructure underwent cellular solidification similar to the AlSi10Mg. Delamination appears to be the dominant failure mechanism for both the X-up and Z-up Charpy specimens. SEM images of the Z-up fracture surface can be seen in Figure 19.Figure 18. SEM images of Al 6061-T6 X-Up Charpy fracture surfaceFigure 19. SEM images of Al 6061-T6 Z-Up Charpy fracture surface3. Conclusions and RecommendationsThe AlSi10Mg samples manufactured, heat-treated and tested in this work displayed relatively isotropic properties.

Conclusions and RecommendationsThe AlSi10Mg samples manufactured, heat-treated and tested in this work displayed relatively isotropic properties. The Al 6061 samples, however, were highly anisotropic and behaved similarly to fibrous composites. This behavior is not surprising considered the tendency of Al 6061 to hot tear during traditional casting and welding processes. A second build ofAl 6061 with pre-heat laser scanning produced parts with significantly alleviated cracking, however more work needs to be performed to completely eliminate cracking concerns.As a first effort to eliminate the cracking issues seen in Al 6061, a microstructure sample design of experiments should be performed with varying laser power and speed to determine whether an optimal range of scanning parameters exists.

If optimizing a single pass is insufficient to alleviate cracking concerns, subsequent optimization of a pre-heat scan could be performed as well.Another potential method to alleviate cracking is raising the platform temperature analogous to increasing the mold temperature in castings. Raising the platform temperature reduces thermal gradients during solidification, reducing thermally-induced stresses. Furthermore, increasing the platform temperature slows the rate of cooling, thereby reducing the microsegregation associated with non-equilibrium solidification. This could be a benefit if liquation cracking was an issue for the Al 6061 samples discussed in this report. The following graph presents the benefit of increasing mold temperature for an aluminum alloy.Figure 20. Benefit of raising mold temperature in casting of Al alloys [3]References[1] J. Davis, Metals Handbook Desk Edition, ASM International, 1998.[2] D.

Benefit of raising mold temperature in casting of Al alloys [3]References[1] J. Davis, Metals Handbook Desk Edition, ASM International, 1998.[2] D. Apelian, Aluminum Cast Alloys: Enabling Tools for Improved Performance, Wheeling,IL: North American Die Casting Association, 2009.[3] S. Li, "Hot Tearing in Cast Aluminum Alloys: Measures and Effects on Process Variables,"2010.[4] "Aluminium Brazing News, Knowledge & Technology," 29 September 2010. [Online]. Available: http://www.aluminium-brazing.com/2010/09/. [Accessed 2 August 2014].[5] J. A. Schey, Introduction to Manufacturing Processes, Third ed., McGraw-Hill HigherEducation, 2000, p. 155..

COMPARISON OF BINDER BURNOUT AND MECHANICAL CHARACTERISTICS OF PRINTED AND CHEMICALLY BONDED SAND MOLDSDean Snelling1, Christopher Williams1, Alan Druschitz2 1 Design, Research, and Education for Additive Manufacturing Systems Laboratory; Department of Mechanical Engineering 2 Department of Materials Science and Engineering Virginia Tech, Blacksburg, VAABSTRACTVarious material systems have been created for Binder Jetting of sand molds; however, a formal analysis comparing the materials to commonly used foundry molding materials has not been conducted. In this paper the authors investigate potential differences in the material properties from four different commercially available binders systems for chemically bonded sand molds. Specifically, the authors compared the binder burnout characteristics and the tensile strength of sand created by 3D printing and conventional chemically bonded molding materials.

Specifically, the authors compared the binder burnout characteristics and the tensile strength of sand created by 3D printing and conventional chemically bonded molding materials. Increased binder content can strengthen the mold but have adverse effect on part quality. Understanding the binder characteristics of printed molds are essential due to the potential defects from large amounts of gas generated from binder while pouring molten metal.1. INTRODUCTION1.1.Direct Printing of Sand Casting Molds via Additive ManufacturingMetal casting is commonly used in manufacturing to create high-quality, low-cost products. While capable of producing a wide range of geometries, the sand casting process inherently constrains the design of a part to geometries that can be produced by forming sand around pre- made patterns and removing the pattern from the sand.

Several “design for manufacturing” guidelines must be followed for successful part design for sand casting, including draft angles for pattern release, and pattern geometries that are able to be removed from sand, etc.Recently, engineers have looked to using Additive Manufacturing (AM) to overcome these geometric restrictions; by processing the sand in an additive context, one is able to directly “print” complex sand molds without the need for a pattern. This pattern-less sand molding approach enables the sand casting of complex mold geometries that could not be made by traditional means thanks to the ability to directly print entire molds (including vents, downsprues, runners, etc.) from a single CAD file (Figure 1) [1,2].

The integration of AM technologies to the foundry industry has increased the efficiency of design and casting production including reducing time and cost associated with the traditional methods of development [3], while still maintaining high accuracy [4].Figure 1. Complex Casting Fabricated via Patternless Sand Molds [2]For example, in Laser Sintering (LS), printed sand molds are created by scanning the laser across a sand bed to selectively melt a resin coated quartz sand mixture in a layer-by-layer fashion [3,5]. An augmented stereolithography process (Optomec), which selectively cures viscous ceramic-resin slurry via exposure to UV energy, has also been used to directly fabricate sand molds [6].The most common AM process for the direct fabrication of sand molds is the Binder Jetting process (Figure 2). In this process, layers of spreadable powdered material is bonded together, via the selective inkjetting of a binder solution into a powder bed.

In this process, layers of spreadable powdered material is bonded together, via the selective inkjetting of a binder solution into a powder bed. A layer of material is spread from the feed tray to the build tray. Binder is jetted over the sand, the build tray lowers, and a new layer is spread. The process is continued layer-by-layer until the part is completed. Figure 2 provides an overview of the Binder Jetting process.Figure 2. Binder Jetting Printing ProcessResearch on Binder Jetting patternless sand molds has been conducted and used for many sand casting applications. Kobliska and coauthors (2005) embedded thin film thermocouples in an A356 matrix cast in molds fabricated by ZCorporation printer [7]. Gill & Kaplas (2009) compared castings printed with ZCast® and Investment casting using starch and plaster, including dimensional tolerances, hardness values, surface roughness, production cost, and shrinkage [8].

Additionally, other studies have been conducted in which it was proved cost effective for use of rapid tooling and prototyping [9].Although the binder jetting process has been used for many applications, including ceramics and powdered metallurgy, much effort has been made for the production and characterization ofpatternless sand molds. However limited research has been conducted to characterize the binder effects and binder burnout in patternless molds produced by Binder Jetting.1.2.Effects of Binder in Sand Molds on CastingsIt is important to understand the strength comparisons and binder burnout characteristics of printed molds as these properties could greatly affect the final part quality of the castings. For example, Meisel, Williams, and Druschitz were unsuccessful in their initial attempts to cast complex cellular structures using 3D printed molds, due to incomplete filling of the mold at the bottom of the structure [1].

This was caused by gas generation from when binders were exposed to the high temperatures of molten metal. Additional research was conducted by Snelling and coauthors (2012) to mitigate these defects in molds produced by a ZCorporation system. The authors conducted casting data, binder burnout percentages, and total binder content from printed molds [10]. It was concluded the ZCorporation system used large amounts of binder to bind proprietary powders, which resulted in the large amount of off-gassing producing incomplete castings. A curing cycle established to decrease the amount of gas generation while maintaining sufficient strength while increasing the quality of castings. Similar tests have been performed by McKenna et al. (2008) to increase the quality of mold by studying the effects of curing time and temperature on permeability and compressive strength [3].

Similar tests have been performed by McKenna et al. (2008) to increase the quality of mold by studying the effects of curing time and temperature on permeability and compressive strength [3]. A mathematical model was used to determine an optimal curing time and temperature for both permeability and compressive strength. It was determined that the best permeability occurred when the molds were baked at 227 °C for 6.2 hours and the best compressive strength at 173 °C for 5.5 hours.Due to the increased binder requirement for the ZCorporation system and a need for thermal curing cycle studies, the authors look to different technologies to form patternless molds for metal casting. ExOne utilizes a two part binder (furan with binder catalyst) that is similar to chemically bonded molding methods in traditional sand casting applications.

ExOne utilizes a two part binder (furan with binder catalyst) that is similar to chemically bonded molding methods in traditional sand casting applications. This binder selection decreases the amount of binder required to create a mold and therefore eliminates the need for post-processing the printed mold in a thermal curing cycle. There exist no prior literature on the binder burnout and gas generation of ExOne system.In this work, the ExOne Sand Materialization System will be compared to four different commercially available traditional chemically bonded sand binders on a basis of binder burnout and strength. The characteristics of binder burnout in printed molds enables an understanding of binder burnout percentage at multiple temperatures.

The characteristics of binder burnout in printed molds enables an understanding of binder burnout percentage at multiple temperatures. This data is then compared to acceptable binder standards and the user is able to select binder type to match casting media (aluminum alloys, iron alloys, steel alloys, etc.).The experimental methods employed in this work, which are focused in comparing the strength and binder burnout of printed molds against traditional molds, are presented in Section 2. Experimental results are presented in Section 3. Conclusions are offered in Section 4.2. EXPERIMENTAL METHODSThe purpose of this paper is to provide a comparison of the ExOne Sand Materialization System to traditional binder systems on the basis of strength and binder burnout.

EXPERIMENTAL METHODSThe purpose of this paper is to provide a comparison of the ExOne Sand Materialization System to traditional binder systems on the basis of strength and binder burnout. This work is guided by the following research question:Research QuestionHow does the printed ExOne two part binder system and sand compare with common chemically bonded sand binders?The ExOne sand material system is compared to four non-printed sands, as listed in Table 1. The sands are compared on the basis of (i) their formed strength and (ii) in the amount of binder burnout. Because chemically bonded sands do not require heat to catalyze a reaction for bonding, two-part polymer binders need to be mixed to create a bond between sand particles. Typically, the binder system consists of a resin and catalyst.

Typically, the binder system consists of a resin and catalyst. There are many resins available including phenolic, furan, alkyds, and sodium silicates, while catalysts include sulfonic acids, isocyanate, and glycerin.The systems used for comparison were chosen because of they were readily available commercially applied binders and catalysts. The Super Set 942 and TW30 was chosen specifically because of the similarity to the ExOne printed resin system (furan resin and binder catalyst).Table 1.

Printed and Non-Printed Binder Systems for ComparisonBinder System ExOne Sand MaterializationAlkyd 18-415 and Alkyd 23-217 Coreactant Super Set 942 and TW30 Chem Bond 490 and Chem Bond 260 Phenoset-RB Part 1 and APR-015 Part 2Printed/Non-Printed PrintedNon-PrintedNon-Printed Non-PrintedNon-PrintedMaterial Type Furan Resin with Binder Catalyst Alkyd Resin and CatalystFuran Resin and Acid Catalyst Sodium Silicate Resin with Glycerin Catalyst Phenolic Resin with Acetate Ester CatalystBecause ExOne systems use a furan binder with a binder catalyst, it is expected to have comparable strength and binder burnout of traditional furan and binder catalyst sands, as presented in the Hypothesis.Research HypothesisPrinted ExOne sands using furan and binder catalyst will compare to traditional sands that use furan and binder catalysts.Although the printed sand and chemically bonded sand both use a furan and binder catalyst will compare in strength and binder burnout percentages with time and temperature, they may slightly differ.

Printed sands need to be well controlled for accurate placement of binders for increased geometrical accuracy. Because of this need, environmental control including humidity and temperature are kept constant to minimize variability. Additionally, the sand distribution iswell controlled (smaller grain size and distribution) which could contribute to differences in properties between the printed and non-printed furan binder systems.2.1.Production of Specimens for ComparisonIn order to compare both printed and non-printed systems, equivalent specimens are needed as a basis for comparison. For evaluation of the sand molds’ strength, an AFS standard 3342-00- S tensile specimen is used as the pattern for comparison [11]. Small cubes of the sands were used for evaluating the sands’ binder burnout characteristics.

Small cubes of the sands were used for evaluating the sands’ binder burnout characteristics. Non-printed sand specimens were formed into standard AFS tensile specimens and then cut into small cubes; printed ExOne rectangles were cut into small cubes.2.1.1.Printed SpecimenPrinted specimens were designed in CAD and printed using an ExOne System to thedimensions given by the AFS Standard (Figure 3). A printed specimen can be seen in Figure 4.Figure 3. AFS Standard Tensile Specimen AFS 3342-00-S [11]Figure 4. ExOne Printed AFS Tensile Specimen2.1.2.Traditional Chemically Bonded SpecimensThe sands were combined with each of the binders using a mixer. Sufficient silica sand is used for each binder system in order to fill the AFS 12-gang tensile specimen box. The resin and catalyst are then measured, added, and mixed according to its vendor’s recommendations, as showing in Table 2.Table 2.

Binder Systems Used for Comparison for Burnout Characteristics [10,12]Binder Addition Resin/CatalystAmountAddition OrderMixing Type1234Alkyd 18-415Alkyd 23-217 CoreactantSuper Set 942TW30Chem Bond 490Chem Bond 260Phenoset-RB Part 1APR-015 Part 2ResinCatalystResinCatalystResinCatalystResinCatalyst1.5 % by weight of sand 20 % by weight of resin 1 % by weight of sand 35 % by weight of resin 3 % by weight of sand 10 % base on weight of resin 4:1 Resin to Catalyst; 1.6 % Binder by weight of sand122121Simultaneous MixtureStand MixerStand MixerStand MixerStand MixerStand MixerStand MixerContinuous Sand MixerImmediately after mixing the polymer binders with silica sand, the mixture is placed into the AFS standard multi-cavity core box to create the tensile specimens for mechanical and binder burnout tests. An image of the multi-cavity core box can be seen in Figure 5.

An image of the multi-cavity core box can be seen in Figure 5. The tensile specimens are dimensionally equivalent to the printed mold (Figure 3).Figure 5. AFS Standard Multi-cavity Core Box2.2.Evaluation of Mechanical Strength of Tensile SpecimenTensile tests are performed to compare strengths of the traditional binders to the printed ExOne binders. Six specimens were tested for evaluation of strength. All specimens are allowed to cure for a minimum of 24 hours to ensure full polymeric cross-linking. The specimens are tested using the AFS tensile specimen fixture (Figure 4) at room temperature on a Com-Ten 95 Series tensile machine at a rate of 0.6 mm/min – 1.2 mm/min.2.3.Binder Burnout CharacteristicsTo evaluate the burnout characteristics of the sand specimens, small pieces of sand bonded with commercially available binders were cut from the remaining tensile specimens (Figure 6). Printed specimens were cut into small cubes from rectangular parts.

Printed specimens were cut into small cubes from rectangular parts. Each material was separated into small crucibles and an initial weight was taken in order to establish a datum. To maintain consistency, approximately 27 g of each materials was measured in each crucible. One-hour curing cycles were performed for all materials starting at 105C to remove all humidity. The temperature was then increased to 150C, to 600C (via 50C increments), and then to 900 C (via 100C increments); each cycle lasted one hour. After each cycle, the cubes’ mass was evaluated via a balance. The mass was divided by the cubes’ initial mass to determine the percent binder burnout.Figure 6. Specimens Used for Binder Burnout Tests3. RESULTS3.1. Mechanical StrengthThe mechanical strength of the printed and non-printed dogbones, as evaluated by tensiletest, are listed in Table 3.Table 3.

RESULTS3.1. Mechanical StrengthThe mechanical strength of the printed and non-printed dogbones, as evaluated by tensiletest, are listed in Table 3.Table 3. Results of Strength Tests of both Printed and Non-Printed AFS DogbonesBinder TypeExOne – Printed Alkyd 18-415 and Alkyd 23-217 Coreactant Super Set 942 and TW30 Chem Bond 490 and Chem Bond 260 Phenoset-RB Part 1 and APR-015 Part 2Mean Strength (kPa)1324.931 964.901 397.028 128.511 561.509 [10]Standard Deviation (kPa) 168.2576 83.6231 46.2907 20.6333 92.5800 [10]From the results, the ExOne printed tensile specimens are significantly stronger than the non- printed binder systems.

The combination of furan binder and binder catalyst and refined sand distribution provide enhanced strength over the common phenolic binders (Phenoset resin and APR catalyst).It was hypothesized that the furan resin with binder catalyst (Super Set 942 and TW30) would have similar characteristics to the printed specimens because they used a similar two part system with minor changes in properties due to better control of sand distributions in the ExOne system. However, this factor exhibits a larger effect than first hypothesized. The ExOne printed specimens maintain sufficient strength for handling and casting alloys.3.2.Binder Burnout CharacteristicsIn order to equally compare binder burnout characteristics, the values for burnout are normalized. This results in a percentage of burnout at each temperature up to 900 °C.

This results in a percentage of burnout at each temperature up to 900 °C. The normalized percentage loss is illustrated in Figure 7.PhenolicSodium SilicateExOne Printed02040608010012002004006008001000Normalized Weight Loss (%)Temperature ( C)Normalized Binder Burnout vs. TemperatureFuranAlkydFigure 7. Weight Loss Comparison of Printed and Non-printed Chemically Bonded SandsFrom the results of the weight loss comparison, there are significant differences in the binder burnout characteristics of all binders in both printed and non-printed systems. For all binders, there is a large jump in burnout at the initial temperature of 105 °C due to the loss in moisture trapped in the specimens. The Chem Bond 490 and Chem Bond 260, as well as Phenoset-RB and APR-015, plateau after the initial loss in moisture. Each binder system contains a temperature where they begin to rapidly decrease in binder content. These are illustrated in Table 4.Table 4.

Each binder system contains a temperature where they begin to rapidly decrease in binder content. These are illustrated in Table 4.Table 4. Temperature of Increased Rate of Binder BurnoutBinder System ExOne – Printed Alkyd 18-415 and Alkyd 23-217 Coreactant Super Set 942 and TW30 Chem Bond 490 and Chem Bond 260 Phenoset-RB Part 1 and APR-015 Part 2 [10]Increased Rate of Binder Burnout 300 °C 250 °C 300 °C 450 °C 150 °CBoth the values of temperature for increased rate of burnout and values at maximum burnout are important to understanding which binder systems are suited for ranges metal alloy. For example, binder burnout at a lower temperature may be more applicable for alloys with lower melting points.

For example, binder burnout at a lower temperature may be more applicable for alloys with lower melting points. Although the binders may burnout at a lower temperature, they may have more desirable features such as lower health hazard, minimal binder requirement, lower cure time, decreased strip time, etc.As discussed in Section 1.1, more binder present increased off-gassing when molten metal comes in contact with the mold, therefore a decrease in quality of the final part. Figure 8 illustrates the weight loss percentage versus temperature. Although all of the specimens completely burnout binder content, the ExOne printed specimens show a binder burnout at much higher temperatures as compared to the other binder systems.

Although all of the specimens completely burnout binder content, the ExOne printed specimens show a binder burnout at much higher temperatures as compared to the other binder systems. This will result in increased quality as compared to other non-printed binder systems (Alkyd 18-415 and Alkyd 23-217 Coreactant, Super Set 942 and TW30, Chem Bond 490 and Chem Bond 260, and Phenoset-RB Part 1 and APR-015 Part 2) as well as other studied printed binder systems for patternless molds (ZCorporation [10]).00.20.40.60.811.21.41.61.802004006008001000Weight Loss (%)Temperature (°C)Binder Burnout vs. TemperatureFuranExOne (Printed)PhenolicAlkydSodium SilicateFigure 8. Total Binder Burnout Weight Loss of Printed and Non-Printed SystemsAdditionally, it was hypothesized that the printed ExOne binder system would behave similar to that of the Super Set 942 and TW30 as both are furan and binder catalyst systems.

From Figure 7 and Figure 8, both the ExOne and Super Set 942 and TW30 binder systems follow a similar trend as expected. However, printed molds held shape at higher temperatures when compared those of the traditional furan (Super Set 942 and TW30) system. Figure 9 illustrates both the ExOne printed mold and Super Set 942 and TW30 systems after binder burnout at 450 °C.a)b)Figure 9. a) Printed ExOne Mold and b) Super Set 942 and TW30 after Binder Burnout at 450 °C4. CONCLUSIONSThe purpose of this paper is to compare specimens on the basis of strength and binder burnout from both printed and non-printed molding techniques. ExOne Sand Materialization System is compared to four commonly used chemically bonded binder systems as a basis for comparison. Understanding these characteristics of printed molds are of increased importance due to the potential defects from large amounts of binder causing excessive amounts of gas in castings.

Understanding these characteristics of printed molds are of increased importance due to the potential defects from large amounts of binder causing excessive amounts of gas in castings. Binder gives strength to the part to maintain structurally sound mold for handling and decrease erosion from molten metal, but consequently generates large amounts of gas producing defects.It was hypothesized that the ExOne specimens would have similar characteristics to the Super Set system as both are furan resin and binder catalyst systems. However, the strength was significantly higher for the ExOne printed sands due the well-controlled sand and binder distributions. The ExOne printed tensile specimens were also much significantly stronger than all other commercially available binder systems tested Table 3.

The ExOne printed tensile specimens were also much significantly stronger than all other commercially available binder systems tested Table 3. The printed specimens were followed similar trends (Figure 7 and Figure 8) in the binder burnout percentages at temperature to the Super Set 942 and TW30, although the ExOne specimens maintained shape longer (Figure 9). The ExOne specimens also had a higher burnout temperature as compared to the remaining binder systems.Overall, when compared to traditional sand, the printed binder system had superior mechanical performance and contained low binder content for gas generation while maintaining its shape at higher temperatures.5. REFERENCES[1] Meisel, N. A., Williams, C. B., and Druschitz, A., 2012, “Lightweight Metal CellularStructures via Indirect 3D Printing and Casting,” SFF Symposium.[2]Snelling, D. A., Blount, H., Forman, C., Ramsburg, K., Wentzel, A., Williams, C. B., and Druschitz, A.

A new, similar process to FDM is the Arburg Plastic Freeforming (APF). The main differences between both processes are the form of the starting material (FDM: Filaments, APF: Conventional granulate) and the material deposition during the layer formation (FDM: Melt strand, APF: fine molten droplets).Since the two processes can be used in similar applications, the aim of this study is to compare both processes in a holistic way. Furthermore, the advantages and disadvantages of the processes are to be highlighted. The systematic comparison between a Stratasys 400mc and the Freeformer 200-3X is divided into the areas of component properties, design limitations and economic efficiency. The material ABS-M30 (Stratasys) is used in both processes.

The material ABS-M30 (Stratasys) is used in both processes. The results show comparable component properties regarding mechanical and optical properties but also differences in design limitations and cost efficiency.IntroductionThe Arburg Plastic Freeforming (APF) is a relatively new additive manufactur ing process with its official market launch in 2013. With this process it is possible to manufac ture three-dimensional components using standard granulates without the need of molding tools [1]. The unique technique of the material deposition in the APF process is opening a new field of possibilities and also new boundary conditions. In contrast to the widely known Fused Deposition Modeling (FDM) process the material in the APF process is not deposited as a strand but in the form of fine molten thermoplastic droplets. As the droplets in the APF process are significantly smaller than the standard strand widths in the FDM process e.g.

As the droplets in the APF process are significantly smaller than the standard strand widths in the FDM process e.g. more detailed component geometries can be realized. Both the APF and the FDM process can be used for similar applications. Until now there is no in-depth comparison of the two processes regarding the component properties, design limitations and economic efficiency.The aim of this research is the development of an overview of the capabilities of both, the APF and FDM process. For this purpose, the same specimens for the evaluation of the different aspects will be manufactured with both processes.

For this purpose, the same specimens for the evaluation of the different aspects will be manufactured with both processes. On both machines, the Stratasys Fortus 400mc (FDM) and the Freeformer 200-3X (APF) the material ABS-M30 will be used to assure comparability while maintaining highest filament quality.State of the ArtThe Arburg Plastic Freeforming is characterized in particular by the processing of standard plastic granulates as well as by the production of components out of very fine molten thermoplastic droplets. The associated machine system for this technology is the Freeformer from Arburg GmbH & Co KG. Its most important machine components are shown in Figure 1.575The raw material, a qualified standard thermoplastic granulate, is fed via a hopper. In the material preparation unit, the granulate is molten with a screw as in the injection molding process. The molten material is then pressed into the material reservoir.

In the material preparation unit, the granulate is molten with a screw as in the injection molding process. The molten material is then pressed into the material reservoir. Here, a piezo actuator performs a pulsed nozzle closure using a needle to close the nozzle opening. The needle in the nozzle moves up and down, producing almost 250 droplets per second. The movement of the building platform, for the precise positioning of the discharged droplets in the x- and y-direction, is realized by two linear motors. After the completion of a layer the building platform is lowered by one-layer thickness, using a spindle drive [1, 2, 3].Figure 1: Schematic setup of the Freeformer [1].The Fused Deposition Modeling process (FDM) was first commercialized in 1991 by Stratasys. As shown in Figure 2 a thermoplastic filament is fed into the heated nozzle where it is molten and pushed through the nozzle opening by the trailing filament.

As shown in Figure 2 a thermoplastic filament is fed into the heated nozzle where it is molten and pushed through the nozzle opening by the trailing filament. Through the defined movement of the FDM head in the x-y-plane the material discharged from the nozzle is deposited in a specific manner. After the completion of the current layer either the FDM head is raised or the building platform is lowered, depending on the machine, by one layer height. Afterwards the deposition process starts again until the desired component is completed [4].Figure 2: Schematic of the FDM process.576A first comparison of the mechanical properties resulting from the FDM and the APF process was conducted by PINTER et al. [5]. In the investigations, components manufactured in the FDM, APF and injection molding process were compared.

[5]. In the investigations, components manufactured in the FDM, APF and injection molding process were compared. Considered properties are the tensile strength, Youngs’s modulus, elongation at break, flexural strength, flexural modulus, flexural strain, impact energy and density.Experimental InvestigationsFor the investigations of the process limit for the FDM and APF process the process limits to be considered are divided into two categories. These are the resolution and the specific limits of the material extrusion. The specimen geometries for each of the categories are given in Figure 3 and are assigned to the specific aspects via the letters a-i.into the following aspects: minimum wall thickness (a), minimum slot width (b), minimum pin diameter (c), minimum hole diameter (d) and minimum part thickness in z-direction (e). The specimen geometries (a) - (d) are in accordance of the standard DIN ISO 52902.

The specimen geometries (a) - (d) are in accordance of the standard DIN ISO 52902. Specimen geometry (e) is according to an investigation carried out by ADAM [6].The investigation of the resolutionis further dividedThe specific limits of the material extrusion are too further divided into the maximum bridging length without support (f), the maximum overhang without support (g) and the minimal angle between building platform and component without support (h). Another aspect of the specific limits of the material extrusion is the achievable length to diameter ratio in z- direction (i). The specimen geometries (f) and (g) are according to FERNANDEZ et al. [7] while (h) is according to the standard DIN ISO 52902. Finally, specimen geometry is according to ADAM.Figure 3: Display of the used specimen geometries.577All investigations are carried out with the ABS-M30 material distributed by Stratasys.

Finally, specimen geometry is according to ADAM.Figure 3: Display of the used specimen geometries.577All investigations are carried out with the ABS-M30 material distributed by Stratasys. This is for assuring the comparability of the specimens between the FDM and APF process while also maintaining a high filament quality. As the slice heights are preselected for differe nt nozzle sizes in the slicing software (Insight) for the Fortus 400mc the layer height for the APF process was adapted accordingly. For a layer height of 0.1778 mm in the FDM process a corresponding layer height of 0.178 mm was set for the APF process. Other relevant process parameters are given in Table 1.

For the tensile test specimens orientated in z-direction in the APF process a special support structure has to be applied to stabilize the part and prevent it from oscillating with raising build height (see Figure 4 right). A soluble support material is used to avoid damaging the tensile test specimens during the removal of the support structures. The specimens manufactured with the FDM process are without a support structure.Figure 4: Orientation of the test specimens (left) and the special support structure for the z-orientation (right).578Specimens for the comparison of the impact strength are manufactured in the same orientations according to DIN EN ISO 179-2, while specimens for the flexural strength are manufactured according to DIN EN ISO 178. All specimens are tested according to the individually applicable standards.Furthermore, specimens for the examination of geometrical properties are built with the FDM and the APF process.

All specimens are tested according to the individually applicable standards.Furthermore, specimens for the examination of geometrical properties are built with the FDM and the APF process. The specimen for the geometrical properties consists out of a variety of different design elements (Figure 5). For example, curved surfaces, holes and circles. With this specimen a direct comparison of several design elements between the two processes is possible. The densities are determined using cubes with an edge length of 20 mm. The assessment of the economic efficiency is described in the separate section “Economic Efficiency” after the evaluation of the experimental results.Figure 5: Specimen for the geometrical properties.Results and DiscussionBeginning with the comparison of the achievable resolution with the FDM and APF process the results for the minimum wall thickness and minimum slot width show that in both aspects both processes deliver similar results (Figure 6).

With both processes only the 0.2 mm and the 0.1 mm thick walls could not be manufactured. It is noticeable that in the FDM process only the wall with a thickness of 1 mm is being filled. All other walls are only consisting out of two contour strands. In contrast the APF process can realize a filling of all manufacturable wall thicknesses (see Figure 6 bottom). This is because the drops have a smaller width than the deposited strands in the FDM process.APF1 mm0.4 mm0.8 mm 0.6 mm0.2 mm 0.1 mm1 mm 0.8 mm 0.6 mm 0.4 mm 0.2 mm 0.1 mmFDMFigure 6: Comparison of the minimum wall thickness (left) with a detail view of the wall thickness of 0.8 mm (bottom) and the minimum slot width (right).579The next two aspects in the evaluation of the experimental results are the minimum pin diameter and minimum hole diameter. Here, pins with the diameters 4 mm, 3 mm, 2 mm, 1 mm and 0.5 mm were manufactured with both processes.

Here, pins with the diameters 4 mm, 3 mm, 2 mm, 1 mm and 0.5 mm were manufactured with both processes. As shown in Figure 7 (left) all pins except for the one with 0.5 mm diameter are manufacturable. Regarding the minimum hole diameter all holes could be manufactured with the APF process. Holes with a diameter of 0.7 mm, 0.6 mm and 0.5 mm could not be manufactured with the FDM process (see Figure 7 right). The reason for this is again the lower drop width in the APF process compared to the strand width in the FDM process.APF0.5 mm4 mm 3 mm 2 mm 1 mm 0.5 mm0.6 mm0.7 mmFDM0.8 mm1 mmAPF FDMFigure 7: Comparison of the pin diameters (left) and the hole diameters (right).The last aspect regarding the resolution of both processes is the minimum part thickness in z-direction. The results show that for the FDM process a minimum part thickness in z-direction of 0.6 mm is necessary to achieve a visually closed part surface.

The results show that for the FDM process a minimum part thickness in z-direction of 0.6 mm is necessary to achieve a visually closed part surface. Using the APF process a part thickness in z-direction of 0.3 mm is sufficient to obtain a visually closed part surface.Following, the results for the specific limits of the material extrusion are presented. The specimens for the maximum bridging length without support show better results in the FDM process. Bridging lengths of up to 10 mm are manufacturable without fatal defects in the specimens. In the APF process a length of 4 mm can be bridged without significant errors. This is because each droplet is deposited individually in the APF process and is only slightly connected to the previously deposited droplet.

This is because each droplet is deposited individually in the APF process and is only slightly connected to the previously deposited droplet. In contrast to that the strand deposited in the FDM process can be stretched over 10 mm without the strand sagging to a critical extend.The specimens for the maximum overhang show no significant differences between the FDM and APF process. Overhangs of up to 0.8 mm can be manufactured with minor defects in both processes. The minimum angle between build platform and component on the other hand shows significant differences between the two processes. In the FDM process the minimum angle where parts can be manufactured without defects is 40°. With 60° the APF process needs significantly higher angles to assure a reliable process. The comparison of the maximum length to diameter ratio (LDR) shows that in the FDM process specimens with a height of 150 mm and a diameter of 5 mm can be realized resulting in an LDR of 30.

The comparison of the maximum length to diameter ratio (LDR) shows that in the FDM process specimens with a height of 150 mm and a diameter of 5 mm can be realized resulting in an LDR of 30. In contrast to that the APF process is only capable of producing specimens with a height of 75 mm with a diameter of 5 mm resulting in an LDR of 15. This is because in the APF process the build platform is moved instead of the head in the FDM process. This causes an oscillation of the specimens at a much lower height in the APF process and leads to lower achievable LDRs. Results for the density show no relevant differences. The specimen for the geometrical properties (Figure 5) shows identical results for both processes.580The results of the determination of the mechanical properties show for most parts simila r values. For the tensile strength, Young’s modulus and elongation at break the values for the xy- and xz-orientation are approximately in the same range.

For the tensile strength, Young’s modulus and elongation at break the values for the xy- and xz-orientation are approximately in the same range. Only the results for the z-orientatio n show significant differences (Figure 8). The reason for this is that probably the heat energy brought into the component by the new deposited material is higher for the FDM process as the strand has a greater volume as the droplet in the APF process. Therefore, the weld seams between the layers are not as distinct in the APF process as they are in the FDM process which leads to lower values for the z-direction.Figure 8: Results of the tensile tests.The results for the flexural tests are shown in Figure 9. The values for the flexuralstrength and modulus are in the same range for both processes in all orientations. Only flexural strain shows noticeable differences between the two considered processes.

The values for the flexuralstrength and modulus are in the same range for both processes in all orientations. Only flexural strain shows noticeable differences between the two considered processes. Here, the APF always delivers higher values, except for the z-orientation where the values are again on a comparable level. The values of the impact strength test show slightly higher values for the APF process (Figure 10).581Figure 9: Results of the flexural tests.Figure 10: Results of the impact strength tests.Economic EfficiencyFor the evaluation of the economic efficiency over 60 influencing factors are considered. In this paper, in total six influencing factors with the highest impact are selected and investigated (Table 2). As standard granulates can be used in the APF process the material costs are significantly lower. As nozzles e.g. are more expensive in the APF process the costs of wear are higher compared to the FDM process.

As nozzles e.g. are more expensive in the APF process the costs of wear are higher compared to the FDM process. The same is observable for the depreciation as the Freeformer that the Fortus 400mc but needs less energy. Also, the maintenance costs are lower in the APF process as they must not be performed by a service technician. The staff costs for the APF process are higher because preparing a build job requires flushing the material reservoir and calibrating the droplet size, which both take time. All in all the resulting costs per hour of build time are on the same level.

All in all the resulting costs per hour of build time are on the same level. But due to the overall higher build rate in the FDM process the FDM process is cheaper.is more expensive582Table 2: Comparison of the most influential costs.APFFDMMaterial costs (ABS)12.50 €/kg216.26 €/kgWear0.09 €/cm30.01 €/cm3Depreciation (5 Years)30,000.00 €/a24,000.00 €/aEnergy961.71 €/a1,361.36 €/aMaintenance3,750.00 €/a8,000.00 €/aStaff costs0,75 h/job with 60 €/h  45 €/job0,58 h/job with 60 €/h  34.80 €/jobConclusion and OutlookThe investigations show that although a different type of material deposition is used in the APF and FDM process the resulting part properties are mostly very close to each other. In the resolution aspect the APF has a slight advantage when it comes to small hole diameters or the complete filling of thin walls.

In the resolution aspect the APF has a slight advantage when it comes to small hole diameters or the complete filling of thin walls. On the other hand, the FDM process is capable of manufacturing lower angles between building platform and component than in the APF process. Also the bridging of gaps of up to 10 mm without support structures is possible in the FDM process. Regarding the mechanical properties the main differences can be observed in z-orientation. Here, the FDM process shows higher values compared to the APF process. The comparison of the economic efficiency shows that the FDM process is more economic efficie nt. Especially when switching to an open parameter FDM machine, which enables the use of cheaper materials, the advantage of the FDM process in regard to the economic efficiency is significant.In further investigations the development to speed up the building process in the APF to catch up to the FDM process is planned.

In addition to that, the economic efficiency shall be further investigated to get an even more detailed insight into the applicable costs. Finally, more process parameters shall be considered for the comparison of the mechanical properties.AcknowledgementThe presented results were obtained in the context of the research project MO 685/19-1 “Investigation and modeling of design and manufacturing guidelines for the plastic freeforming by use of ABS considering the process-specific material degradation” sponsored by the Deutsche Forschungsgemeinschaft (DFG). We would like to thank the DFG for its support.References[1]H. Gaub: “Customization of mass-produced parts by combining injection molding and additive manufacturing with Industry 4.0 technologies”, Reinforced Plastics, Volume 60, Number 6, 2016.583[2] M. Neff, O. Kessling: “Geschichtete Funktionsteileimindustriellen Maßstab“,Kunststoffe, Nr. 08/2014, pp. 64-67, 2014.[3]O.

mparison of Compression Molding and Selective Laser Sintering Processes in the Development of Composite Bipolar Plates for Proton Exchange Membrane Fuel CellsEhsan Taghipour, Ming C. Leu, Nannan Guo, Department of Mechanical and Aerospace Engineering Missouri University of Science and Technology, Rolla, MO 65409AbstractBipolar plates are key components of Proton Exchange Membrane (PEM) fuel cells. They carry current away from the cell and withstand the clamping force of the stack assembly. Therefore, PEM fuel cell bipolar plates must have high electrical conductivity and adequate mechanical strength, in addition to being light weight and low cost in terms of both applicable materials and production methods. In order to attain these goals, we have manufactured graphite-carbon- polymer composite plates using Compression Molding (CM), which is suitable for mass production, and Selective Laser Sintering (SLS), which is suitable for making prototypes.

In this paper, the electrical conductivity and flexural strength of the bipolar plates fabricated using the CM process versus constitutive materials are experimentally studied. The properties of bipolar plates fabricated using the CM process are compared with those of plates fabricated using the SLS process. Natural graphite (NG), synthetic graphite (SG), carbon black (CB), and carbon fiber (CF) are used as the constitutive materials for both processes, with epoxy resin employed as the binder matrix. By varying the volume fraction of each constituent, the distribution of the electrical conductivity and flexural strength of parts made using the CM and SLS processes are obtained, and the similarities and differences of the effects of the various constituents between these two processes are compared.Keywords: Bipolar plates; Graphite-carbon-polymer composite Compression molding; Selective Laser Sintering.1.

IntroductionProton exchange membrane (PEM) fuel cells are attractive options for alternative energy production to fossil fuels due to promising features that allow them to generate electrical energy without undergoing combustion, such as their high power density, relatively low operating temperature, good start-up, ability to convert fuel to water as the only byproduct, longer lifetime, and shape modularity [1,2,3].

In a fuel cell stack, bipolar plates are key components as they serve the following functions [3,4,5]: (i) (ii) (iii) (iv) (v)distribute the fuel and oxidant within the cell, facilitate water management within the cell, separate the individual cells in the stack, carry current away from the cell, enable heat transfer,226(vi) (vii) withstand the clamping forces of the stack assembly, and (viii) prevent leakage of reactant gases and coolant.support thin membranes and electrodes,The bipolar plate is the heaviest component of the PEM fuel cell, accounting for about 80% of the total weight of a fuel cell stack. It is also a very costly component. According to a cost analysis study, about 45% of a fuel cell stack’s cost is incurred by the bipolar plate [1]. Therefore, quality improvement and weight and cost reductions are critical issues in the development of PEM fuel cell bipolar plates, especially those fuel cells employed in automotive or mobile applications.

The processing methods and materials used to manufacture bipolar plates determine the final properties of the products, so they are of great importance for any manufacturer or researcher of the bipolar plates.Materials that have been used for making PEM fuel cell bipolar plates include non-porous graphite (both natural and synthetic), metals (both non-coated and coated), and polymer composites [1]. Graphite is the most commonly used material for fabricating PEM fuel cell bipolar plates because it has a suitable combination of thermal and electrical properties and excellent corrosion resistance in the highly-acidic PEM fuel cell environment. However, graphite is very brittle and has poor mechanical properties, machining graphite plates is very costly and time-consuming. Metallic bipolar plates are stronger and more conductive, and also can be fabricated more easily and at a lower cost than graphite bipolar plates.

Metallic bipolar plates are stronger and more conductive, and also can be fabricated more easily and at a lower cost than graphite bipolar plates. However, they suffer from low corrosion resistance in the PEM fuel cell environment. Graphite-carbon- polymer composites are promising alternatives to both graphite and metal because they have the advantages of good mechanical, thermal, and electrical properties; they are also light weight and have high corrosion resistance [6,7]. Conductive materials such as flaky natural graphite, synthetic graphite, expanded graphite, and carbon black have been used by many researchers to impart good electrical conductivity to composite bipolar plates [3,5,6,8-14]. Moreover, carbon fiber has been employed to improve the mechanical integrity of bipolar plates.

Moreover, carbon fiber has been employed to improve the mechanical integrity of bipolar plates. The Department of Energy (DOE) proposed a technical target for bipolar plates for the year 2010 [15], in which the main requirements were to achieve electrical conductivity greater than 100 S/cm and flexural strength greater than 25 MPa.In addition to materials, the manufacturing method and processing conditions of composites greatly influence the dispersion and inter-particle positions within the polymer matrix; therefore, they have significant effects on the final electrical and mechanical properties of the composite plates [7]. Compression molding (CM) is a common method for making polymer composite bipolar plates [16] and is very suitable for mass production. The literature reveals that it has been used widely for making graphite-carbon-polymer composites, either with thermosetting or thermoplastic resins.

The literature reveals that it has been used widely for making graphite-carbon-polymer composites, either with thermosetting or thermoplastic resins. The effects of conductive and reinforcing particles on bipolar plate properties including electrical conductivity and flexural strength have been studied experimentally by various researchers [3,5,6,8-13].Selective Laser Sintering (SLS), an additive manufacturing technique, has been studied in making graphite composite bipolar plates for PEM fuel cells [14,17-19]. In the SLS process, a mixture of graphite material and binder in a powder bed is scanned by laser, and the molten binder bonds graphite particles together to form 3D parts layer by layer.

In the SLS process, a mixture of graphite material and binder in a powder bed is scanned by laser, and the molten binder bonds graphite particles together to form 3D parts layer by layer. The major advantages227of the SLS process include its ability to build complex flow fields for bipolar plates and its consumption of less time and financial resources when making bipolar plates from each design, which reduces the time and cost of fabricating bipolar plates in the R&D stage as compared to conventional methods.In the study described in the present paper, the effects of constitutive materials, including natural graphite (NG), synthetic graphite (SG), carbon black (CB), and carbon fiber (CF), on the electrical conductivity and mechanical strength of composite bipolar plates for PEM fuel cells are experimentally studied.

The electrical conductivity and mechanical strength results obtained using the CM process are compared with the results previously obtained using the SLS process [14]. Variations in the electrical and mechanical properties versus changes in the bipolar plate constituents are investigated for the two processes.2. Materials and ToolsThe materials and compositions used for making composite bipolar plates using the CM process are chosen to match those used previously in the SLS process [14] in order to allow more meaningful comparisons between the two processes.2.1 Conductive and reinforcing constitutive materialsNatural graphite (3610), synthetic graphite (4437), carbon black (5303) and carbon fiber (AGM99) are obtained from Asbury Graphite Mills, Inc. (New Jersey, USA). Properties of these materials, as provided by the manufacturer, are given in Table 1. Natural graphite is the primary electrical conducting material, but it has poor wettability with the liquid resin.

Natural graphite is the primary electrical conducting material, but it has poor wettability with the liquid resin. Synthetic graphite has lower electrical conductivity than natural graphite due to its lower degree of crystallinity. Carbon black, manufactured by the combustion or thermal decomposition of hydrocarbon fuel under reducing conditions, has a very small particle size and thus a very large specific surface area. Carbon fiber is used to enhance the mechanical strength of composite bipolar plates.2.2 Polymer resinThe thermosetting epoxy resin is inexpensive and has good mechanical strength, hardness, and thermal and chemical resistance. Liquid epoxy EPONTM Resin 828, an undiluted, clear, difunctional bisphenol A/epichlorohydrin derived liquid epoxy resin, is used in the experiments.

Liquid epoxy EPONTM Resin 828, an undiluted, clear, difunctional bisphenol A/epichlorohydrin derived liquid epoxy resin, is used in the experiments. The density of this epoxy resin is 1.16 g/cm3, and EPIKURETM 3230 is a good hardener for EPONTM 828 when employed in the compression molding processes, according to the information provided by the manufacturer (MOMENTIVE Inc.).2.3. Hot pressA 250 kN hydraulic press with hot platens is used to compress the composite powder in a circular mold. A temperature control unit on the apparatus digitally controls the temperature of the hot platens. Fig. 1(a) shows the hydraulic press with the heated platens between which the mold and dies are located, and Fig. 1(b) shows the dies.2282.4. Steel moldTwo cylindrical steel molds (of 3-inch diameter) and cylindrical punches are used to create the samples.

1(b) shows the dies.2282.4. Steel moldTwo cylindrical steel molds (of 3-inch diameter) and cylindrical punches are used to create the samples. With a band saw, the fabricated composite plates are cut into several specimens with dimensions according to appropriate ASTM standards, which will be described later.Table 1 Properties of graphite materials.PropertiesParticle size (µm) Theoretical density (g/cm3) Surface Area (m2/g) Typical Conductivity (S/cm)Natural Graphite (3610)75-1502.26 1.27 27.78Synthetic Graphite (4437)10-452.26 11.46 17.24Carbon Black (5303)<0.031.8 254 2.93Carbon Fiber (AGM99) Diameter 7.4 Length 150 1.75 1.87 -abFig. 1 (a): The 250 kN hydraulic press with heated platens, and (b): dies used in the experiments.3. Manufacturing Process and Property CharacterizationTo reduce its viscosity, the epoxy resin is kept at 100 oC for several hours.

Manufacturing Process and Property CharacterizationTo reduce its viscosity, the epoxy resin is kept at 100 oC for several hours. The mold, which first is preheated until it reaches 200 °C, is thermally isolated inside the space between the two platens of the hot press to maintain this temperature with the help of thermally insulating fibers. A thermocouple is used to check the temperature of the mold before pouring the powder into it. Next, the hardening agent is added to the epoxy resin according to the prescribed volume percentage, and they are mixed until the liquid becomes clear. Because a 35 vol.% binder was used in all of the SLS experiments, a 35 vol.% resin is used in all of the CM experiments. The epoxy resin and hardener are added to an acetone solvent in a beaker and stirred by a mechanical stirrer for several229minutes.

The epoxy resin and hardener are added to an acetone solvent in a beaker and stirred by a mechanical stirrer for several229minutes. Next, the reinforcing constituents, i.e., natural graphite, synthetic graphite, carbon fiber, and carbon black, are added to the solution according to the specific composition data. The mixture is stirred for over one hour until the acetone evaporates. Then, the beaker is kept under a fume hood for several hours to evaporate the remaining acetone completely. Finally, the mixture is maintained inside an oven at 100 oC for half an hour to make sure that the graphite-carbon-polymer composite powder is completely dry and ready to be compression molded. Compression molding is performed by applying 54 MPa of pressure within 30 minutes of the molding time.The four-probe technique is employed for measuring electrical conductivity according to the ASTM C611 standard.

The KEITHLEY 220 programmable current source is used to apply a constant current of 10 mA between the two outermost current probes, and the KEITHLEY 181 NANOVOLTAMETER is used to measure the dropped voltage within a 20 mm distance between the two innermost voltage probes. Six specimens are selected for data measurement, and the average values are reported for the conductivity. The three-point bending test is performed to measure flexural strength. In accordance with ASTM D790-10, four tests are conducted using the INSTRON 4469 testing machine. The microstructures of the samples are obtained using the Hitachi S-4700 FE-SEM.4. Results and DiscussionThe influences of the constitutive materials on the electrical conductivity and flexural strength of the composite bipolar plates are presented and discussed below.4.1 Effect of synthetic graphiteThe microstructures of parts with variations in natural graphite (NG) and synthetic graphite (SG) are shown in Fig.

2, in which the larger flaky particles are NG and the smaller spherical ones are SG. Electrical conductivity and flexural strength variations with different volume fractions of SG are shown in Figs. 3(a) and 3(b), respectively. As the volume fraction of SG increases, the conductivity of the bipolar plate decreases for both the CM and SLS processes because the conductivity of SG particles is lower than that of NG. The current conduction among NG particles is reduced by the presence of SG particles when the two types of particles are mixed together. Increasing the SG content increases the flexural strength of the samples made using the CM process but decreases that of the samples made using the SLS process. In the CM process, the stress distribution around small spherical SG particles is more uniform and the strain is smaller, so the mechanical strength improves.

In the CM process, the stress distribution around small spherical SG particles is more uniform and the strain is smaller, so the mechanical strength improves. On the other hand, in the SLS method, adding smaller SG particles fills up the big pores among larger flaky NG particles, which reduces the porosity, so less resin can infiltrate, leading to lower mechanical strength.230(a)(b)Fig. 2 SEM images showing the microstructures of (a): a brown part made using the SLS process and (b): a part made using the CM process with 15 vol.% SG, 50 vol.% NG and 35 vol.% binder.(a)(b)Fig. 3 Variations in (a) electrical conductivity, and (b) flexural strength versus changes in the volume fraction of SG for the CM and SLS processes.231The experimental results indicate that the electrical conductivity and flexural strength obtained when using the CM process are much lower than the SLS results and are below the DOE target values.

The maximum electrical conductivity value with the CM process is approximately 74 S/cm, while the maximum value with SLS is approximately 380 S/cm. The low electrical conductivities in the CM parts have been observed by previous researchers. Dweiri and Sahari [9] used synthetic graphite, carbon black, and polypropylene with the melt compounding preparation method. Their composite bipolar plates fabricated using the CM process had a maximum electrical conductivity of 36 S/cm. Lee et al. [8] also developed compression molded bipolar plates using powder epoxy resin, graphite, and carbon black, obtaining an electrical conductivity below the DOE target.

Lee et al. [8] also developed compression molded bipolar plates using powder epoxy resin, graphite, and carbon black, obtaining an electrical conductivity below the DOE target. Bhlapibul and Pruksathorn [20] achieved a maximum electrical conductivity of only 4.52 S/cm when they employed polyester resin and graphite powder to fabricate bipolar plates using the CM process.The measured electrical conductivity and flexural strength in the bipolar plates made using the CM and SLS processes can be attributed to the materials and preparation methods employed in these processes. In the CM process, the liquid resin covers the surface of the graphite particles during mixing and thus hinders the electrical connection between the graphite particles. In the SLS process, the phenolic binder in the form of particles is mixed with the graphite particles, and the polymer resin is infiltrated into the brown part, imparting mechanical strength to the composite plate.

The phenolic binder does not remain inside the part after sintering; instead, it is converted to carbon ashes during sintering. The carbon ashes fill up the pores and help to increase the electrical conductivity. On the other hand, in the CM process, the mixture does not contain sufficient epoxy resin to impart mechanical strength to the fabricated parts. Moreover, there is local mixing of the liquid epoxy resin with the composite plate particles (rather than with the overall powder constituents uniformly), so agglomerates form inside the mixture after the acetone solvent is incorporated and the mixture is dried because of the accumulation of epoxy resin in some regions. As a result, the electrical conductivity and flexural strength decrease as the agglomerates introduce inhomogeneity inside the structure.

As a result, the electrical conductivity and flexural strength decrease as the agglomerates introduce inhomogeneity inside the structure. Based on the previous research results of CM manufactured bipolar plates with high electrical conductivity [3,5,6,7], mixing ‘solid’ epoxy resin with the filler powder or using graphite particles with longer aspect ratios may help resolve the problem of low electrical conductivity and flexural strength of CM manufactured bipolar plates.5.2 Effect of carbon fiberCarbon fiber is used widely in composite materials to increase mechanical strength. Fig. 4 shows the microstructures of a sample made using the CM process and a brown part made using the SLS process, with CF and NG as their constituents, in which long, thin carbon fibers are uniformly mixed with natural graphite particles.Variations in electrical conductivity and flexural strength versus changes in the CF volume fractions are shown in Figs. 5(a) and 5(b), respectively.

5(a) and 5(b), respectively. Both CM and SLS processes show the same trend in electrical conductivity and flexural strength. Adding CF into NG reduces the electrical conductivity because the dispersion of CF among NG232particles lessens the contacts of NG particles, while CF itself has much less conductivity. Increasing the CF ratio increases the flexural strength of the composite samples, as expected. In Fig. 5 the electrical and mechanical properties of the CM fabricated parts using these NG, CF, and epoxy resin materials are lower than those of the SLS fabricated parts and the DOE targets. This is due to the lack of homogeneity in the mixture of materials and the formation of polymer agglomerates inside the plate structure during preparation for the CM process, as explained before.(a)(b)Fig.

4 SEM images showing the microstructures of (a): a brown part made using the SLS process, and (b): a part made using the CM process with 25 vol.% CF, 40 vol.% NG and 35 vol.% binder.(a)(b)Fig. 5 Variations in (a) electrical conductivity, and (b) flexural strength versus changes in the volume fraction of CF for the CM and SLS processes.2335.3 Effect of Carbon BlackFig. 6 shows the microstructure of a brown part made using SLS and a part made using CM, both with 16 vol.% CB, 49 vol.% NG and 35 vol.% binder. The surface of NG particles is covered by the nano-size CB particles in the SLS sample shown in Fig. 6(b); however, CB particles are present between NG particles but do not cover the surface of NG particles in the CM sample shown in Fig. 6(c).Fig. 7 shows variations in electrical conductivity and flexural strength with different volume fractions of CB.

6(c).Fig. 7 shows variations in electrical conductivity and flexural strength with different volume fractions of CB. In the CM process, adding CB to NG decreases the electrical conductivity because CB has less intrinsic electrical conductivity than NG. However, when the CB volume fraction continues to increase, the electrical conductivity increases because nano-size CB particles fill the voids between NG particles. Likewise, increase in CB can enhance the flexural strength after a certain volume fraction. In the SLS process, as the CB content increases, electric conductivity decreases. This is because CB has lower intrinsic conductivity than NG. During the ball-milling process, CB particles with large surface areas tend to agglomerate and cover the whole surface of NG particles, as shown in Fig. 6(b), which impedes contact between NG particles.

6(b), which impedes contact between NG particles. This is different from the CM process, in which CB particles fill the small voids between NG particles rather than covering their surface; thus, electrical conduction can pass through these small CB particles in CM made bipolar plates.234(c)Fig. 6 SEM images showing (a) microstructure of a brown part made using SLS; (b) detailed view of the surface of NG particle; (c) microstructure of a part made using CM (16 vol.% CB, 49 vol.% NG and 35 vol.% binder).(a)(b)Fig. 7 Variations in (a) electrical conductivity, and (b) flexural strength versus changes in the volume fraction of CB for the CM and SLS processes.2355. ConclusionThe compression molding (CM) and selective laser sintering (SLS) processes are compared in terms of their fabrication of composite bipolar plates for PEM fuel cells.

ConclusionThe compression molding (CM) and selective laser sintering (SLS) processes are compared in terms of their fabrication of composite bipolar plates for PEM fuel cells. The effects of these two processing methods on the electrical conductivity and flexural strength of the graphite-carbon- polymer composite bipolar plates are evaluated. For purpose of meaningful comparison, the same materials and compositions are used in both manufacturing processes. Different volume fractions of NG, SG, CF, and CB, with a constant amount of epoxy resin as the binder, are used for comparison in the experimental study. The results show that the trends of electrical conductivity variations versus changes in the volume fractions of SG and CF are the same for the CM and SLS processes; however, increasing the volume fraction of nano-size CB increases electrical conductivity in the CM process but decreases electrical conductivity in the SLS process.

mparison of density of stainless steel 316L parts produced with selective laser melting using different powder grades A. B. Spierings*, G. Levy*institute for rapid product development irpd, inspire AG, Lerchenfeldstrasse 5, CH-9014 St.Gallen Abstract:Selective Laser Melting is a powder based additive manufacturing process where the metallic powder particles are fused to 3D parts using a high energy laser beam. Much work has already been conducted to investigate the details of the process, suitable materials and process parameters and further more. As metallic powders are the raw material for this process, there are still a lot of open questions relating to suitable grain size distributions for dense parts with regard to productivity, surface quality, mechanical strength and ductility. The present work shows the results of density measurements of parts, produced using three different particle size distributions and different energy densities of the laser beam.

The present work shows the results of density measurements of parts, produced using three different particle size distributions and different energy densities of the laser beam. Two layer thicknesses of 30μm and 45μm were investigated. It is shown that without a minimal amount of fine grains, which are able to fill the voids between the coarse grains, lower scan speeds are needed in order to produce dense parts. Furthermore, the differences in the relation of the powders to the densities, the layer thicknesses and laser scan speeds indicate, that the powder grain size distribution plays an important role and that should be taken into account for optimal results. This work is a contribution to the ASTM initiative F42 for “Additive Manufacturing”.I.

This work is a contribution to the ASTM initiative F42 for “Additive Manufacturing”.I. Introduction Looking back to the development directions of conventional production technologies like drilling and milling, rapid technologies designate a logical development stage towards further integration of computer technologies into the production processes [1]. In the very beginning of industrial production, extensive manual labour was predominant for hundreds of years. The invention and integration of electric motors into the production processes led to a huge increase of productivity, accuracy and repeatability. The next important step was the invention of electric control systems, which further improved the reliability of the processes. Nevertheless, manual work was and is still important, although many preparation tasks for production can be done at the desktop and the production processes are usually fully computer controlled.

Nevertheless, manual work was and is still important, although many preparation tasks for production can be done at the desktop and the production processes are usually fully computer controlled. The next step towards a further integration of computer techniques into the production processes was laid by the additive manufacturing processes. These technologies need only a CAD data set of the item to build. Therefore, no further programming steps e.g. of the production machines have to be undertaken – it is a fully integrated process from data generation (CAD) over process planning, parameterisation to the self running production [2]. As a consequence, by reducing manual efforts to a minimum, these techniques have the potential for the next step towards higher productivity.

As a consequence, by reducing manual efforts to a minimum, these techniques have the potential for the next step towards higher productivity. The group of corresponding additive production processes [3] include Selective Laser Sintering and Selective Laser- or Electron Beam Melting; so powder bed based processes to create plastic or metallic parts, respectively. Other similar processes are for example stereo lithography and 3D printing, which uses a fluid resin or a photopolymer, respectively, to build up parts. Today, additive manufacturing processes like e.g. SLS and 3D Printing, tend towards Rapid Manufacturing (RM), where the parts produced are used directly in the end application and not only as a prototype. The SLM process would be perfectly usable for RM, as the metallic parts produced provide high material strength [3, 4] due to the achievable density of nearly 100% bySFF Symposium 2009Review PaperA. Spierings, inspire - irpdfully melting the particles [4].

Spierings, inspire - irpdfully melting the particles [4]. Furthermore, the industry requires metallic materials for RM [5]. Although SLM parts can be used in a wide field of applications like medical implants [6-10], mechanical engineering [11] and aerospace industry [12], there are still many challenges to be solved for a real rapid manufacturing. Beside others, these challenges include the part orientation, layering strategy, support generation and structure, process speed and control facilities [5], which have to be considered in the future [13]. By improving these subjects, additive processes for producing metallic parts (SLM, EBM) could be some when considered as real RM-processes in the future. Over all, today additive processes can be considered as entering the standardisation stage [3, 14] as first RM applications have already been conducted successfully [15, 16]. This is underlined by first efforts for an international standardisation by ISO [17, 18] and ASTM.

This is underlined by first efforts for an international standardisation by ISO [17, 18] and ASTM. The later has started the initiative F42 for Additive Manufacturing Processes in 2008. This paper is a contribution to open questions in the field of Selective Laser Melting. Even for stainless steel 316L, there are a number of different sources available providing powers with different particles size distributions. Furthermore, the layer thicknesses of the parts to be produced vary between 30μm and about 60μm, depending on the application, the geometry, the machine used and the preferences of the operator. The paper presents results concerning differences in achievable densities using three different particles size distributions using layer thicknesses of 30μm and 45μm. The results contribute also important information to current standardisation initiatives (ASTM).II.

The results contribute also important information to current standardisation initiatives (ASTM).II. Methods, Materials and Measurementsa) SLM Machine Metallic parts have been produced using a SLM machine of Concept Laser GmbH, Germany, type M1 (2005). This machine is equipped with a Q-switched Nd: YAG diode-pumped solid state laser source of Rofin-Sinar Laser GmbH. This laser provides a continuous laser beam with a wavelength of 1064 nm, the spot size is 0.2mm and the nominal laser power at the building platform is in the range up to about 104W ± 0.5W. The machine software uses a specific scan strategy where the cross sections of the geometries are divided into 5mm x 5mm squares. Each square is shifted from layer to layer by 1 mm in x- and y- direction. Furthermore, the squares are rotated by 45° to the x- or the y-direction of the building platform.

Each square is shifted from layer to layer by 1 mm in x- and y- direction. Furthermore, the squares are rotated by 45° to the x- or the y-direction of the building platform. These arrangements allow it to minimise any isotropy in x- and y-direction.b) Materials Each test cube was produced in 316L stainless steel, whereas three different particles size distributions were used (Table 1). The grain size distributions were measured using a HELOS&RODOS R3 Laser Diffraction Sensor of Sympatec GmbH.Powder No. 316L – Type 1316L – Type 2316L – Type 3D10 [μm] D50 [μm] D90 [μm] D50 / D10 Tap density [g/cm3] Relative powder density Span S [ - ] [19]6.3 15.05 30.79 2.4 4.8 ≈ 60.1% 1.6319.91 28.19 41.31 1.4 4.6 ≈ 57.6% 0.7615.64 37.00 59.69 2.4 4.7 ≈ 58.9% 1.19Table 1: 316L powders used for producing the test cubes.31.08.2009SFF Symposium 2009Review PaperA.

Spierings, inspire - irpdThe powders are characterised in the following way: Powder type 1 and 2 (Figure 3) can be characterised by a typical Gaussian like grain size distribution with a slight asymmetry towards coarser grains where as type 3 (Figure 3) is an asymmetric distribution with an increased concentration of finer grains. Furthermore, the particles are not as spherical as e.g. type 1 and in comparison the maximal grain sizes are much bigger. Although type 1 has a significant higher amount of fine particles, the flowability of the powder, which has not been measured, is still acceptable.Figure 1: SEM picture of powder type 1 (left) and type 3 (right)Figure 2a: Size frequency and cumulative size frequency of powder type 131.08.2009SFF Symposium 2009Review PaperA.

These parameters can be divided into four groups: Material specific parameters (grain size distribution, flow properties etc), laser parameters (laser power, spot size etc.), scan parameters (scan velocity, hatching distance, etc.) and environmental parameters (ambient temperature, protective gas atmosphere, O2 level). These four groups of parameters are listed in more detail in [20]. The current machine setup allows only the adaption of scan speed, laser power, hatching distance (Table 2), and some more, less important parameters. The cubes produced have a dimension of 12x12x12mm3 and were produced with the processing parameters according to Table 2. The ambient temperature was approximately 40°C as there is no preheating of the powder or the process chamber.

The ambient temperature was approximately 40°C as there is no preheating of the powder or the process chamber. The O2-level during the SLM process was between 0.3% and 0.8%.Materials 316L – Type 1 30 104 300 – 8002 0.13316L – Type 2 30 104 250 - 500 0.13Layer thickness tLayer [μm] Laser Power1 PLaser [W] Scanning speed2 vScan [mm/s] Hatch distance sHatch [mm] Table 2: Processing parameters for the production of test cubes 1 Laser power approximately 104W ± 0.5W 2 The scan speed was adjusted in steps of 25 mm/s or 50 mm/s316L – Type 3 30 104 250 - 500 0.13A layer thicknesses tLayer of 30μm was used as this thickness typically leads to good the surface qualities of the parts produced and stair effects are minimal [20]. The laser power PLaser was set to the maximum possible power in order to maximise the scan velocity and productivity, respectively.

The laser power PLaser was set to the maximum possible power in order to maximise the scan velocity and productivity, respectively. Expecting better density values for powder Type 1 compared to the coarser powders, the scan speed range was set somewhat higher than for powder types 2 and 3. A driving parameter for the density of SLM-parts is the energy density Ed [20], which includes all above mentioned process parameters:P sLaserE=(1)d⋅ With PLaser the laser power, vscan the scan velocity, sHatch the hatching distance and tLayer the layer thickness. TSpot [19] describes the time dependency of the melting process and related effects like e.g.

TSpot [19] describes the time dependency of the melting process and related effects like e.g. the temperature dependence of the melt viscosity, with ds the spot size diameter:v⋅tscanHatchLayerTspot=d vSScan(2)d) Density measurement for bulk test cubes The density of a SLM-part was measured using the Archimedes method [21], which allows to measure the mean density of any complex geometry, as not only single cross-sections are taken into account but the whole volume of a part. For conducting the measurements, a Mettler balance (Type AE200) with a specific density measurement device for solid materials (Type AB33360) was used. The calculation of the density ρp of the part under consideration follows equation (1):31.08.2009SFF Symposium 2009Review PaperA.

The calculation of the density ρp of the part under consideration follows equation (1):31.08.2009SFF Symposium 2009Review PaperA. Spierings, inspire - irpdρ p=m a mm −afl⋅ρ fl(3)ρfl is the density of the fluid, which is temperature dependent, ma is the mass of the part in air and mfl is the mass of the part in the fluid. The measurements were conducted with de-ionised water. The temperature-dependence of the fluid [22] was taken into account.III. Results and Discussiona) Considerations to powder layer density For Additive Manufacturing Powder Processes such as SLS and SLM, the powder layer density should be as high as possible [19] in order to produce dense parts with high scan velocities and therefore with high productivity. The density of a powder layer is mainly dependent on the particle sizes or size distributions, respectively. If the amount of fine particles, e.g.

The density of a powder layer is mainly dependent on the particle sizes or size distributions, respectively. If the amount of fine particles, e.g. in the range of about 5μm or less, is too high, the agglomeration of particles can eliminate their positive effects of filling up voids. McGeary [23] showed in his experiments that a bi-modal mixture of powders with a size ration of 1:7 can increase the powder density by about 30%, if the amount of fine particles reaches 30%. Regarding the density of thin powder layers, as used for SLS and SLM, the situation is different compared to typical density measurements of powders (apparent and tap density) [19]. The reason is that the particle diameters typically can be in the same range as the powder layer thicknesses (e.g. 30μm and 45μm as in this study). Karapatis [19] showed that the additional amount of voids, caused by the platform, can cause as much as 40% of the total porosity of the powder bed.

30μm and 45μm as in this study). Karapatis [19] showed that the additional amount of voids, caused by the platform, can cause as much as 40% of the total porosity of the powder bed. This amount tends to 0 the thicker the layer is. Therefore, it is generally helpful to use particle size distributions where the fine particles can fill the voids between the bigger particles. Furthermore, spherical particles are preferred for high packing density since they exhibit low interparticle friction and high mobility [24]. The measurement of the real density of a thin powder layer is very difficult. Karapatis [19] described in his experiments that the layer can be compacted by the deposition system used and that the layer density can be increased to values higher than apparent or tap densities.

Karapatis [19] described in his experiments that the layer can be compacted by the deposition system used and that the layer density can be increased to values higher than apparent or tap densities. An optimum ratio of 10 particles per layer thickness leads to a layer density of more than 60%, which is slightly more than the measured tap density of the powders used in this study.Real powder layer thickness teff Assuming a powder layer density of 60% (Table 1) within the first generated 30μm layer, the scanning to a material density of e.g. 99% leads by shrinkage in the vertical direction to a scanned layer thickness of 18.2μm. Together with lowering the building platform by 30μm, the resulting next powder layer thickness is 41.8μm. After building less than 10 layers (Figure 4), the real, by the deposition system generated powder layer thickness reaches a stable value of 49.5μm (for 60% powder layer density), which is then scanned to a bulk layer thickness of 30μm.

The optical penetration depth, which describes the depth at which the laser intensity equals the incident intensity be e, is about 63μm, measured for a fine titanium powder [26], indicating that this distance is just in the range or somewhat bigger than the effective layer thickness. Furthermore, within the first about 6 to 10 layers the real thickness of the structure being built is smaller than the theoretical thickness of these first layers. Using the situation above, the difference within the first 7 layers is about 20μm. As this difference is typically within a part- specific support structure, the part itself is not affected by these inhomogeneous layer thicknesses so that its dimensional accuracy in the vertical z-direction is given.b) Part Density In order to get insight into the fluctuation range of density measurements, three samples produced with the same process parameters were measured four times each.

The mean density of these parts was 99.53% and the standard deviation of the density measurements was 0.082%. A repetition was performed using slightly changed process parameters, which resulted in a mean density of 99.33% and a standard deviation of 0.070%. If the amount of energy, which is brought into this powder layer is sufficient high to fully melt all the powder particles, nearly fully dense parts are producible with all three powder types. Due to the different powder characteristics (Table 1), the scan velocities for dense parts can vary significantly.Layer thickness 30μm With 30μm layers it is possible to produce parts with very high densities (> 99%, typically about 99.5%) if enough energy is available to fully melt the powder particles. The lower the energy densities) are the lower are the part densities, as reported by [21]. A further parameter affecting the solidification behaviour is the time constant (equ.

The lower the energy densities) are the lower are the part densities, as reported by [21]. A further parameter affecting the solidification behaviour is the time constant (equ. 2) of the process, which has to be high enough compared to typical time dependent physical effects of a melt pool like e.g. the break up time of a scan track. As soon as tSpot >≈ 0.0035s nearly fully dense parts are possible (Figure 6) with all powder types. This is nearly one order of magnitude longer than the typical break up time of a melt pool [27].31.08.2009SFF Symposium 2009Review PaperA. Spierings, inspire - irpdAs expected, Figure 5 shows that the density is also highly dependent on the powder characteristics. Regarding the 30μm layers, it is obvious that the coarser powders typically result in lower density values at a specific energy density. There are two main factors influencing porosity: First, the energy supplied might not be sufficient to fully melt the bigger particles.

There are two main factors influencing porosity: First, the energy supplied might not be sufficient to fully melt the bigger particles. Therefore, these particles can remain, especially at the lower not irradiated sides, partially unmelted, which generates voids within the scanned layer. A second, indirect effect is the emerging surface roughness of parts, produced with coarse powders. Coarse powder particles are not able to fill the valleys between the scan tracks of the last layer, which generates also voids. So, if the amount of fine particles is big enough, mentioned voids can be partially filled and easily be melted by heat conduction.

Spierings, inspire - irpdLayer thickness 45μm In contrast to the 30μm layers, the densities of the 45μm layers are - for a specific energy density or scan speed, respectively - in the range of about 1% lower for all three powder types (Figure 7). This can be expected and is obvious as the energy flux is reduced quickly the thicker the powder layers are [28]. Interestingly and in contrast to the 30μm layers, powder type 2 shows for all scan speeds a better density than powder type 3. This could be explained by the fact that the effective layer thickness for 45μm layers reaches up to about 74μm (Figure 4). As all particles are smaller than this effective thickness, the amount of fine powder grains becomes comparatively less important. It is more important to consider the amount of bigger particles and their diameters. Regarding this, powder type 3 has the biggest particles, which need more energy to get fully molten.

It is more important to consider the amount of bigger particles and their diameters. Regarding this, powder type 3 has the biggest particles, which need more energy to get fully molten. Therefore, the corresponding part densities are lower than for powders with finer grains. Figure 5 shows that the densities of the 45μm / powder type 2 measurements are almost identical to the 30μm / powder type 1 values.

Figure 5 shows that the densities of the 45μm / powder type 2 measurements are almost identical to the 30μm / powder type 1 values. The better values of a 45μm-layer processed with powder type 1 (Figure 7) is explained by the fact that the comparatively lower scan speeds allow more time for heat conduction and therefore a better melting process of the fine particles.Density stainless steel 316L 45um100.0%316L - Type 3 - 45um99.0%98.0%316L - Type 2 - 45umy t i s n e d97.0%96.0%316L - Type 1 - 45um95.0%94.0%93.0%92.0%20.030.040.050.0 70.0 60.0 Energy density (J/mm3)80.090.0100.0Figure 7: Dependence of the part density of powder type and energy density for 45μm layersc) Discussion The results of this study indicate that it is possible to produce parts with a high density (≈ 99.5%) using all three different grain size distributions, when the scan speeds are adjusted adequately.

If the particle size distribution does not fulfil some basic requirements [24], the adjustments of the process parameters typically affect the productivity of the SLM process in a negative way. One has to be aware that a remaining porosity in the range of about 0.5% can not be avoided in powder bed based processes like Selective Laser Melting or Selective Laser Sintering. There will always be a remaining porosity in the parts produced. The goal is to minimise this remaining porosity, where as the definition of an acceptable porosity is dependent on the process of comparison. Looking at the SLM process and current results, the achievable densities are at least as good or better than the densities of parts produced with other forming production techniques like e.g. Metal Injection Moulding [29, 30] or casting. The porosity of additive manufactured parts is affected by several influencing parameters:31.08.2009SFF Symposium 2009Review PaperA.

Metal Injection Moulding [29, 30] or casting. The porosity of additive manufactured parts is affected by several influencing parameters:31.08.2009SFF Symposium 2009Review PaperA. Spierings, inspire - irpdMaterial type:Different materials, such as steel, Ti, CoCr, and Aluminum highly affect, due to their specific physical behaviour, the details of the melting process and therefore of the binding mechanism [31].

However, the general results and dependencies are valid for all types of metal materials [10, 20, 32].Powder type:(cid:131) Grain size distribution: Relevant powder parameters are at least D10, D50and D90.(cid:131) The powder bed density [19, 24], which is affected by the grain sizedistribution and the coating system (ruler, flexible coating system, roller), should be as high as possible.(cid:131) The effective powder layer thickness (Figure 4).Karapatis [24] derived criteria for suitable particle size distributions, using readily available numbers:D90<tLayer,D D5010≥10,D D9010≤19(4)These equations indicate that 50% of the particles are 10 times coarser than the 10% finer grains and that about 20% of the particles sizes are in a 1:20 ratio. The requirements have been found basically on investigations for bi-modal powder size distributions and adapted for mono-modal powders.

The requirements have been found basically on investigations for bi-modal powder size distributions and adapted for mono-modal powders. However, for the powders used in this study, the second requirement is not fulfilled as this ratio is between 1.4 and 2.4 (Table 1) and the third requirement is not fulfilled either. Furthermore, although there are significant differences between powder Type 1 and Type 3, the D50/D10 – ratio is about the same.IV ConclusionAdditional requirements for thin powder layers have to be defined in order to assess adequately typically available powders for SLM. It is difficult to generate general requirements based on the three available different powder size distributions used in this study. Any requirement should be able to assess the quality of powders considering flowability and resulting part density and productivity (scan speed) in a SLM process. Nevertheless, some basic requirements for powders are presented.

Nevertheless, some basic requirements for powders are presented. It is assumed thatt eff 90D This indicates that the effective powder layer thickness teff is at least 50% higher than the diameter of 90% of the powder particles. Therefore, most of the particles can be deposited within teff.≈ 1.5D D This requirement asks for a sufficient amount of fine particles, which are able to fill the voids between the coarser grains. Both requirements together indicate that D10 is about 7.5 times smaller than teff.90≈ 51031.08.2009SFF Symposium 2009Review PaperA. Spierings, inspire - irpdIn addition to these requirements, which address the part density and the productivity of the process, upper limitations exist for the amount of fine particles. If D10 is below about 5μm to 6 μm, agglomeration of the powder particles starts to occur.

If D10 is below about 5μm to 6 μm, agglomeration of the powder particles starts to occur. This reduces the flowability of the powder so that it becomes more and more difficult to create appropriate, homogeneous powder layers with a high powder layer density [19]. Therefore, Karapatis requires that the Haussner ratio is below 1.25 [19, 33]. However, ρtap and ρapparent are not always available. Further work has to be done in order to define suitable and easily measurable values for different powders.Outlook Besides part density and productivity, the resulting mechanical strength and ductility is also affected by the powder type. Furthermore, it can be assumed that a coarser powder leads to a rougher surface quality, especially when some coarse powder grains are not fully melted by the laser beam.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCOMPARISON OF FATIGUE PERFORMANCE BETWEEN ADDITIVELY MANUFACTURED AND WROUGHT 304L STAINLESS STEEL USING A NOVEL FATIGUE TEST SETUPM. M. Parvez*, Y. Chen*, J. W. Newkirk†, F. F. Liou*Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65401, USA†Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65401, USAAbstractIn this research, a novel adaptive controlled fatigue testing machine was designed for bending type high cycle fatigue test. A unique dual gauge section Krouse type mini specimen was designed for simply supported transverse bending. Displacement controlled fatigue tests were implemented using an electromechanical actuator.

A unique dual gauge section Krouse type mini specimen was designed for simply supported transverse bending. Displacement controlled fatigue tests were implemented using an electromechanical actuator. The variation in the control signal and load observed during the test provides unique insights into realizing the deterioration of the specimen due to fatigue. These analyses were utilized to compare the fatigue performance of wrought and additively manufactured 304L stainless steel. The influence of the build direction on fatigue performance was also investigated by testing specimens with 0, 45, and 90 degrees build direction. These comparisons were carried out at different levels of displacement amplitude.IntroductionAdditive manufacturing (AM) is a process of joining materials into a 3D part layer by layer from 3D computer data, which is opposite to the traditional manufacturing process of using subtractive method [2].

Due to the additive approach, AM has some advantages when compared with subtractive manufacturing. AM process can fabricate complex 3D structures, reduce the lead time, and improve the efficiency of raw materials [3]. During the past decade, AM has been studied in various types of materials, including metals, polymers, glass, ceramics, and composite materials. As metals and their alloys are of great importance in day to day life, AM of metals has become popular in recent years, in which austenitic stainless steel is widely used in various industrial fields. Selective laser melting (SLM) is one of the AM methods used in the field of metal AM [4]. The mechanical properties characterization of austenitic stainless steel using SLM has attracted a lot of interest.

The mechanical properties characterization of austenitic stainless steel using SLM has attracted a lot of interest. Alsalla [5] reported the tensile and yield strength of SLM-processed 316L stainless steel are higher than cast materials, while the elongation is relatively weaker because of the porosity generated due to the gas pore or lack-of-fusion. Yusuf et al. [6] reported the hardness in SLM 316L stainless steel was higher than wrought parts. Casati et al. [7] studied the microstructure of SLM 316L stainless steel. Both columnar and equiaxial grains were formed which is different from traditional 316L stainless steel. Montero [8] investigated the effect of heat treatment on SLM 316L stainless steel. Since most of the mechanical failures of structures are due to fatigue, the fatigue testing study is indispensable in failure mechanism analysis.

Since most of the mechanical failures of structures are due to fatigue, the fatigue testing study is indispensable in failure mechanism analysis. Fatigue properties of several popular metal alloys made by SLM such as Ti-6Al-4V [9], 15-5PH [10], AlSi10Mg [11], and Inconel 718 [12] have been investigated.353In this study, we compared the fatigue performance of additively manufactured 304 stainless steel material built at different orientation with wrought bulk material. A novel displacement-controlled fatigue testing method was implemented to conduct fully reversed (R = - 1) bending fatigue test on mini specimens since axial loading tests on mini specimens suffer buckling. During both monotonic and cyclic loading test, mini specimens are preferred in AM material’s mechanical properties characterization due to the high cost of manufacturing and specimen preparation [13].

But mini specimens exhibit a higher mechanical strength including fatigue compared to the standard specimens [14-19] because standard specimens have greater risk volume to contain larger defects what lead to crack growth resulting in fatigue failure [20-26]. Nonlinear stress distribution in bending type test is another reason for mini specimens to demonstrate higher fatigue strength [27, 28]. But for the comparative study of fatigue performance, mini specimens were used in this study. To obtain a constant stress distribution within the gauge section, a Krouse type specimen with a dual gauge section was designed for simply supported loading. The dual gauge section increases the risk volume and maintains one-dimensional loading. The specimen was designed for simply supported loading because at a given load, the displacement is four times the displacement in a fully clamped loading.

The specimen was designed for simply supported loading because at a given load, the displacement is four times the displacement in a fully clamped loading. We also designed and implemented nonstandard testing equipment for the displacement-controlled test to reduce the cost since standard testing equipment is very expensive [29]. The designed test equipment also consumes less electrical power because of the tests performed on mini specimens require a smaller range of loads.During the displacement-controlled fatigue test, the applied load decreases with the decrease in the stiffness of the material. As the load decreases, the control signal decreases as well. The change in the control signal provides a unique insight into realizing the deterioration of the material in terms of the nucleation and propagation stage. Extended studies have been carried out to develop techniques to identify the nucleation and propagation phase [29-35].

Extended studies have been carried out to develop techniques to identify the nucleation and propagation phase [29-35]. In this research, we introduced a control signal monitoring (CSM) method to identify the phases. This method was applied to compare the fatigue performance of SLM fabricated and wrought 304 SS material.Experimental SetupA dual gauge section Krouse type mini specimen was designed in this research work. Figure 1 illustrates the drawing of the wedge-shaped specimen. The specimen is clamped at both ends with line contacts for simply supported loading. The length of the specimen between both clamping ends is 1 inch, while the total gauge length is approximately one-third of the effective length. The advantage of using dual gauge is the load distributed symmetrically along with the gauges. The fatigue failure can occur randomly at any gauge.

The advantage of using dual gauge is the load distributed symmetrically along with the gauges. The fatigue failure can occur randomly at any gauge. Moreover, the dual gauge section covers a large surface area with increased volume to capture all different types of surface and microstructural defects. The common defects in AM materials are void, pores, microcracks, and lack of fusion. AM materials have a higher probability of defects compared to wrought materials. A sensitivity and uncertainty analysis was also performed to determine the optimal thickness (0.65 mm) of the specimen since mini specimen needs strict dimensional tolerances. In summary, the specimen was specially designed for fatigue testing of additive manufacturing materials. The designed specimen type, size, and loading mechanism add a novelty to the current study.The test materials used in this study are hot rolled and annealed 304 bulk material, and SLM fabricated bars of 304L stainless steel powders.

The chemical composition of both the wrought material and powder is listed in Table 1. From the chemistry of both materials, the354wrought material may be characterized as 304 SS, and the powder as 304L SS due to its low content of interstitial, N, C, O, etc. The relatively close chemistry of materials may add a better understanding of the comparative study though the weight percentage of Ni is 2% higher in the powder but not expected to make a significant difference in test results. In the SLM process, a laser beam was applied on the powder bed of a Renishaw AM250 machine to selectively melt and join a layer of powder particles. Once a raster pattern layer was finished, the next layer was fabricated with 670 rotation angles. Finally, the steel bars were fabricated layer by layer using a set of optimized parameters listed in Table 2 to yield maximum part density.

Finally, the steel bars were fabricated layer by layer using a set of optimized parameters listed in Table 2 to yield maximum part density. Fine surfaced finished (average Ra = 0.482 μm, average Rz = 4.242 μm) specimens were prepared using a wired electric discharge machine (W-EDM); so the effect of surface roughness on initiating the crack would be minimal and the failure would occur only due to the presence of major defects within the volume. A schematic representation of the SLM specimens (not in scale) prepared at different build direction is shown in Figure 2.Mini specimens require a lower range of loads. A low power subwoofer with soft mechanical suspension was used in this study as an actuator since the mathematical modeling of a subwoofer is relatively similar to an electromagnetic actuator. To accommodate the actuator with the test bench, the dust cap of the voice coil was replaced with a lightweight plastic flange.

To accommodate the actuator with the test bench, the dust cap of the voice coil was replaced with a lightweight plastic flange. A load cell installed inline between the central clamp and the flange measures the tensile and compressive force. A laser displacement sensor was fixed with a guide rail to measure the displacement at high speed. Figure 3 illustrates a specimen installed on the fatigue test bench for simply supported loading. The specimen is clamped at the center with a surface contact (3 mm × 7 mm), while at both ends the specimen is in line contact with the bearings. Bearing minimizes frictions during the test. Spacers were accommodated between top and bottom of the bearing adapters at both ends to ensure no preload on the specimen. Toggle clamps were used at both clamping ends. Implementing a microcontroller, the measured displacement data were processed to calculate the mean and amplitude of the displacement. The load values were logged using a data acquisition card.

Implementing a microcontroller, the measured displacement data were processed to calculate the mean and amplitude of the displacement. The load values were logged using a data acquisition card. An adaptive proportional and derivative (PD) controller was implemented to control the amplitude of the displacement. To maintain the desired displacement, the control signal value was sent to a function generator via Ethernet. A linear power amplifier coupled with the function generator drives the actuator. The displacement and control signal amplitudes were stored for the entire fatigue life cycle to monitor the performance of the controller and to analyze the control signal in identifying the nucleation and propagation phases.Results and DiscussionsA fully reversed (R = -1) bending type fatigue test was performed at room temperature on fine finished wrought and AM specimens. The specimens were actuated for closed loop displacement-controlled sinusoidal excitation at 56 Hz test frequency.

The specimens were actuated for closed loop displacement-controlled sinusoidal excitation at 56 Hz test frequency. The occurrence of fatigue failure at a random location within the gauges, as shown in Figure 4, verifies the design effectiveness of the specimen. Figure 5 illustrates the competency of the adaptive PD controller in maintaining the set displacement amplitude. The displacement amplitude goes very high when the final failure occurs. The control signal amplitude decreases until the final failure. Since the stiffness of the specimen reduces as the crack grows, propagates, and final failure occurs, the load to maintain the desired displacement reduces as well.

The stiffness of a material decreases during the nucleation and propagation phase though the rate of the change of the stiffness is different at different phases. Analyzing the control signal, we identified the change in phase from nucleation to propagation. Figure 6 shows the analytical result of the CSM method. Performing a linear regression analysis on the control signal, we estimated the phases. The analysis was applied for all the specimens tested.357To determine the maximum nominal stress on the specimen, the average of the peak load up to the end of nucleation cycle was calculated.

In comparison to the SLM fabricated materials themselves, the specimens prepared at 00 build direction has a higher fatigue strength than the specimens prepared at 450, and 900 direction. The specimens prepared at 00 build direction includes a smaller number of layers at the cross-section than the specimens built in other directions. As a result, the specimens have a lower probability of defects within its volume since a smaller number of layers results in a smaller number of interlayer zones. Interlayer bonding is weak and it has a higher possibility of creating defects during the fabrication of AM materials [30, 31].For further analysis, the end of nucleation cycle and cycles to failure were plotted for different stress levels. The comparison was done between wrought and SLM fabricated specimens built at 00 orientation, and between the SLM fabricated specimens prepared at 450, and 900 build direction (shown in Figure 8).

The comparison was done between wrought and SLM fabricated specimens built at 00 orientation, and between the SLM fabricated specimens prepared at 450, and 900 build direction (shown in Figure 8). At all cases, the number of cycles during both the nucleation and propagation cycles increases as the stress value goes low. This validates the SLM fabricated materials follow the conventional phenomena of nucleation and propagation. At the high cycle fatigue, the propagation cycle for the SLM material is less than the propagation of the wrought material. Authors believe that the presence of defects within the volume of the material enhances the propagation rate.358(cid:3007)(cid:3051)(cid:2870)(cid:3)(cid:4666)(cid:883)(cid:4667)(cid:3)(cid:4666)(cid:884)(cid:4667),...__ 0.300 T l v 0.250 -0 B :..= s' 0.200 - S 0.150alCl(1) u algo 0.100 i50.I.250005000075000100000125000, r 0.7~0.6~0.5 150000(1) "O .e ..... i al - - «i ~ 6b ~ .....

.A : ■.A♦ ♦ .A.A.. ~ ..200l.E+04l.E+05l.E+06 Number of cycles (log scale)l.E+07l.E+08Figure 7: S-N plot of wrought and SLM specimens prepared at different build directionThe loading direction during the fatigue test of the specimens prepared at 00 build orientation is normal to the layer direction, while at 450, and 900 build direction, it is parallel. Therefore, the comparison of the results for 450, and 900 build orientation is pertinent. As shown in Figure 4, the fatigue crack propagates across the cross-section. A crack must travel through a larger number of layers (illustrated in Figure 2) during both the nucleation and propagation for the specimens built at 450 orientation than the specimens at 900 orientation.

Apart from the presence of defects, this is another reason for the higher number of propagation and nucleation cycle for the specimens built at 450 orientation.ConclusionIn this study, the transverse bending fatigue test was performed on Krouse type mini specimen eliminating the buckling issue with miniature specimen occurred in an axial fatigue test. The dual gauge section captures a large extent of microstructural and surface defects increasing the surface area and volume. The influence of build direction on fatigue performance of SLM fabricated 304L stainless steel was compared applying control signal monitoring method. The analysis illustrates SLM materials demonstrate lower fatigue strength at both nucleation and propagation phase compared to wrought materials.

SLM 45 degree: cycles to failure.A. SLM 45 degree: end of nucletion cyclee SLM 90 degree: cycles to failuree SLM 90 degree : end of nucleation cycle... • •• • • ............1-----------l500,0001,000,0001,500,000'--------1...__2,000,000------lI 2,500,000Number of cycles(b)Figure 8: End of nucleation phase and cycles to failure of (a): wrought specimens and SLM fabricated specimens at 00 build orientation, (b): SLM fabricated specimens at 450 and 900 build orientationAcknowledgmentThis research was supported by National Science Foundation Grant CMMI-1625736. Part of the work was also funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.3611.2.3.4.5.6.7.8.9.10.11.12.13.14.15.16.17.ReferencesDiemar, A., R. Thumser, and J.

viewed PaperA COMPARISON OF LAYER DEPOSITION AND OPEN MOLDING OF PETG BY FUSED PELLET FABRICATION IN AN ADDITIVE MANUFACTURING SYSTEMA. Gibson*, J. M. Weaver* *Department of Manufacturing Engineering, Brigham Young University, Provo, UT 84602AbstractAdditive manufacturing continues to offer new possibilities in both production and economics. The industry has quickly adopted it to rapidly produce parts that would be difficult or cost preventative otherwise. Recent innovation has expanded its capabilities, however there are still significant limitations. Most AM processes are restricted by materials available, in producing large parts, or by not achieving material deposition speeds to make certain products feasible. In addition, tight tolerances for features and surfaces cannot be produced without substantial post processing.

In addition, tight tolerances for features and surfaces cannot be produced without substantial post processing. High-speed Fused Pellet Fabrication (FPF) in combination with Hybrid Manufacturing (HM) offers expanded capabilities as additive and subtractive process are used within the same space. It also allows for a different kind of additive process where an open mold can be cut from a substrate and then filled using the FPF process to fabricate parts without layers. This, in combination with Large Area Additive Manufacturing (LAAM), enables parts that leverage the strengths of new and traditional methods at scales and speeds previously unavailable.IntroductionAdditive Manufacturing (AM), often referred to as 3D printing, is a process that creates a three-dimensional part by building it up layer by layer. There are a variety of polymer materials that an AM system can utilize through deposition or curing to create each layer.

There are a variety of polymer materials that an AM system can utilize through deposition or curing to create each layer. These materials are usually in the form of a filament, wire, powder, or resin. Each have their advantages and drawbacks, but they all allow for parts that would be difficult, time consuming, or impossible to create through machining or other traditional subtractive methods. Even still, there are limitations on the speed and scale that these processes can produce at. For decades AM has been relegated to prototyping, this however is changing because of advances in hardware and software. The flexibility and increasing speed allow for smaller runs or one-offs to be produced with lead times and costs that make it feasible and profitable. Multi-material parts are also becoming more common as they offer improved properties and functionality [1].

Multi-material parts are also becoming more common as they offer improved properties and functionality [1]. More and more goods are being produced directly by AM processes and there is increased effort to expand the capabilities of AM technology.High-speed Fused Pellet Fabrication (FPF) systems are a recent innovation in additive manufacturing. This system is a pellet fed which allows for more material to be rapidly extruded than in a filament or powder-based machine. A specially designed extrusion screw retrofitted on to existing spindle-based platforms (such as a CNC mill) heats, stirs, and pressurizes the material and extrudes it as fully melted plastic. Being spindle driven allows for the production of parts in which the size is only limited to by the size of the machine it is attached to. This is one example of what is referred to as Large Area Additive Manufacturing (LAAM) [2,3,4].

This is one example of what is referred to as Large Area Additive Manufacturing (LAAM) [2,3,4]. In addition, a CNC mill brings with it the ability to utilize traditional machine tools in-line with the LAAM process.1011This combination of additive and subtractive capabilities is called hybrid manufacturing and presents great potential for allowing new methods and strategies to rapidly create parts and tooling that would have been more difficult, expensive, or time intensive to produce through traditional methods.Hybrid systems will be able to leverage this in switching between the additive extrusion head and subtractive tools. This would open doors to increased production of multi-material parts and provides more options in design [1]. There is also growing interest in printing parts made from low friction polymers for moving parts, mechanical systems and joints.

There is also growing interest in printing parts made from low friction polymers for moving parts, mechanical systems and joints. There has been much research done for traditional methods, but only on limited materials in additive and not on a hybrid system [5,6]. It is well suited for the hybrid process as friction interactions between bearing surfaces require tighter tolerances then any additive process can deliver alone.MethodologyThe devices used in the project include a HAAS Mill and a Hybrid Manufacturing Technology XTRUDE (FPF additive process) There is a hopper that dries and feeds the pellets into the print head where the mill spindle drives a heated screw which melts, compresses, and extrudes plastic into a continuous 6mm bead.

The coolant lines were modified to blow compressed air in order to regulate the temperature of the material as each layer was printed in order to keep the part wall as uniform as possible and reduce warpage from the heat that would accumulate as the part was built up. In order to produce the most consistent flow of material for testing samples, a simple pattern was selected so that parts would be as consistent as possible in process as well as dimensions. The design was then processed by a developmental beta slicer created by Oak Ridge National Labs [13]. This produced G code that was run on the mill where the feed rate and spindle speed were tuned over a number of experimental runs to get the best results for the individual plastic. In this paper the results from the PETG samples produced will be discussed.To characterize the material properties of the prints, Type 1 samples (165mm overall length) according to the ASTM D638 standard were to be tested for tensile strength.

Because of the nature of layer deposition, and the weakness associated with interlayer bonds, samples were produced in both the in line with direction of deposition (x direction) and perpendicular to the direction of deposition (z direction). The x direction samples were run with the screw temperature of 265 degrees Celsius, spindle speed of 90 rot/min, a feed rate of 0.029 m/s, depositing 196 cubic mm of material per second, and the compressed air blowing at 15 psi. The Z direction prints were run at the same speed and feed, but with 30 psi as the smaller footprint and taller part would accumulate more heat during printing.1012Figure 1: FPF Layer Deposition Printing (X Direction Part)Figure 2: FPF Layer Deposition Printing (Z direction part)Figure 3: FPF Open MoldingIn addition to this, this system can extrude much larger amounts of molten plastic than other forms of additive manufacturing. This allows for a different approach to depositing material to form a part.

This allows for a different approach to depositing material to form a part. In this study it took the form of an open mold. An aluminum mold was created with a slot 32mm in depth, 6mm wide and 178mm long with a 1-degree taper from the base to the opening to allow clearance for the part to be removed after cooling. The print head would then hold 3mm from the opening and extrude into the mold with a spindle rate of 30 rot/min and move across the opening with a feed rate of 0.021 m/s until the cavity was filled. This produced a continuous part that was free of layers.After all parts were produced in each orientation, 5 standard size samples of each were machined where material was removed from all 6 sides using a 12.7 mm 2 flute end mill at a spindle speed of 1500 rot/min and a feed rate of 0.016 m/s. They were then cleaned and deburred1013in preparation of the testing (See Figure 4).

They were then cleaned and deburred1013in preparation of the testing (See Figure 4). The samples were kept in a dehumidifying cabinet after machining to maintain the integrity of the material. These samples were then prepared and tested on a Mini Instron device with an Epsilon laser extensometer, pulled at a rate of 46 mm per minute (See Figure 5).Figure 4: Machined and Prepared SamplesFigure 5: Tensile TestingFigure 6: Tested SamplesResultsTable 1 shows the results for the test samples with the layers in the x direction or direction of deposition (layer direction), and samples of the z direction or perpendicular to the direction of deposition (against layer direction), and the samples that were open molded (no layers).

Yield strength was determined by the 0.2% offset method.Table 1: Tensile Test ResultsAverage Tensile Yield StrengthStandard DeviationLayer direction51.17 MPa0.67133Against Layer Direction47.82 MPa1.13075Open Molded49.41 MPa1.34927Manufacturer Injection Mold Data [11]50 MPan/aSome notable characteristics of each set of samples is that both the open molded and layer direction samples experienced significant elongation, ranging from 6mm to 18mm. There were a few samples the broke after the maximum tensile strength both sets, but the majority had visible plastic deformation between the maximum tensile point and break points. The reverse was true for1014the samples with layers against layer direction. There was one sample that had this significant plastic deformation, the rest however broke cleanly soon after the maximum tensile point.DiscussionSome interesting observations can be made from the data collected.

The difference between the layer deposited samples is apparent, though there is only about 6.5% less strength in the samples against layer direction compared to those in layer direction. The standard deviation between them shows that there is more variance in the performance in the against grain samples, this is expected as the stress is applied in a direction that tests the strength of the interlayer bonds. It is much more likely to have defects and cavitations in the inter-layer bonds and region then within the layer itself. In comparison to a Fused Filament Fabrication (FFF) process, the yield strength is significantly higher. Data from manufactures indicate that the typical yield strength for PETG filament produced parts is about 50 MPa in layer direction and 30 MPa against layer direction [12]. The FPF process can produce much stronger parts.

This is likely the case because the layers are much larger and so the number of layers in each sample is comparatively reduced and the pellets are fully melted in the process so the temperature between the layer deposited and the previous layer is higher, allowing for a stronger bond to form.70.0060.0050.00) a P M40.00(s s e r t S30.0020.0010.000.000.000.100.200.30 Strain0.400.50Sample 1Sample 2Sample 3Sample 4Sample 5Figure 7: Layer Direction Stress-strain10150.60) a P M(s s e r t S) a P M(s s e r t S70.0060.0050.0040.0030.0020.0010.000.0070.0060.0050.0040.0030.0020.0010.000.000.000.00Sample 1Sample 10.100.100.30 Strain Sample 3 Figure 8: Against Layer Direction Stress-strain0.200.40Sample 2Sample 40.200.30 Strain0.40Sample 2Sample 3Sample 4Figure 9: Open Molded Stress-Strain10160.500.50Sample 5Sample 50.600.60The samples that were open molded show a mix of results from the other sample sets.

They are very close only having about 3.5% less strength that the average of the layer direction samples. There is a higher standard deviation in the sample results than the layer deposited samples; this could be that because there are no layers and from the nature of how the plastic flow is more turbulent as it fills the mold, thus the defects and cavitations are more randomized throughout. This demonstrates that the open molding process can produce parts that are high performing, can be produced quickly, and are not as anisotropic in nature in com parison to the layer deposited parts.

This demonstrates that the open molding process can produce parts that are high performing, can be produced quickly, and are not as anisotropic in nature in com parison to the layer deposited parts. The machining of the additive parts to meet tolerance of the ASTM demonstrates that a Hybrid process can indeed produce parts of both high tolerance and of good mechanical properties.Additionally, this study shows that FPF, in combination with hybrid manufacturing, can allow for parts to be produced quickly and cheaply that can perform very close to that of injection molded parts of the same material without the overhead costs of an injection machine and molds. This method spans the gap of purchasing extruded plastic and the time and wasted material of having to machine down from stock where lower part runs are required and the need to produce parts in numbers that justify the cost of injection molding equipment.

Open molding would also allow for shapes and sizes off parts that would be difficult or expensive to produce in small numbers.The next step in future research is to employ this method and process to create multi- material parts. This will be done by filling a cavity in a substrate with the FPF process and then machining the low friction plastic to a specified dimension and surface finish to compare the durability of this multi-material part with one that relies on mechanical fastenings.1017References[1] A. Bandyopadhyay, B. Heer, Additive manufacturing of multi-material structures, Materials Science and Engineering: R: Reports 129 (2018) 1-16. [2] N. Volpato, D. Kretschek, J.A. Foggiatto, C.M. Gomez da Silva Cruz, Experimental analysis of an extrusion system for additive manufacturing based on polymer pellets, The International Journal of Advanced Manufacturing Technology 81(9) (2015) 1519-1531. [3] D. Moreno Nieto, V. Casal López, S.I.

[3] D. Moreno Nieto, V. Casal López, S.I. Molina, Large-format polymeric pellet-based additive manufacturing for the naval industry, Additive Manufacturing 23 (2018) 79-85. [4] E.D. Weflen, A hybrid additive and subtractive manufacturing approach for multi-material components, Iowa State University, 2020. [5] O.A. Mohamed, S.H. Masood, J.L. Bhowmik, A parametric investigation of the friction performance of PC-ABS parts processed by FDM additive manufacturing process, Polymers for Advanced Technologies 28(12) (2017) 1911-1918. [6] I. Maries, C. Vilau, M. Pustan, C. Dudescu, H. Crisan, Determining the tribological properties of different 3D printing filaments, IOP Conference Series: Materials Science and Engineering, IOP Publishing, 2020, p. 012022. [7] R. Walter, R. Selzer, M. Gurka, K.

Multiple laser scans, varying in scanning speed at constant power, were examined before the melting laser scan (pre-heating) or after it (post-heating). The analysis focused on microstructural features such as porosity, and α-phases lath structure, as well as the hardness response of the material. The results revealed the additional layerwise scans had a significant impact on reducing porosity by up to 98% when the additional scan was applied prior to or upon the melting scan. Additional laser scans decreased the microstructure and mechanical response variation along different orientations.

Additional laser scans decreased the microstructure and mechanical response variation along different orientations. Furthermore, these findings highlight the potential of layerwise heating strategies to improve the overall quality and performance of L-PBF Ti-6Al-4V components, thus paving the way for enhanced applications in various industries such as aerospace.Keywords: Laser powder bed fusion, Ti6Al4V, Preheating Scan, Post-heating Scan, Microstructure1.IntroductionTi6Al4V material is one of the most preferred titanium alloys renowned for its outstanding mechanical properties [1]. Its remarkable strength-to-weight ratio, extraordinary fracture toughness, resistance at corrosive environments, and biocompatibility have made it an approved choice for manufacturing durable lightweight components needed for the automotive, aerospace, and biomedical industries [2].

These exceptional aspects have gained Ti6Al4V significant popularity for additive manufacturing (AM) applications [2].AM technology was invented in the 1980s and emerged as a revolutionary process of joining materials layer upon layer to fabricate parts directly from the digital 3D file of the geometry. This technique has shown rapid development due to the superior fabrication ability of complex geometries compared to traditional fabrication techniques. Additionally, the near-net-shape production capability of AM technology enhanced the sustainability of the fabrication by reducing material waste and energy needs required for conventional manufacturing. With the advancement of novel free-form design approaches by topology optimization, AM technology continues to gain momentum by eliminating the design limitations of traditional fabrication methods [3].

Laser powder bed fusion (LPBF) is the most adequate technique for these ingenious design approaches among all the other AM technologies due to its ability to fabricate fine features with high resolution [4].Utilizing the capabilities of the LPBF techniques on Ti6Al4V material allows the fabrication of complex-shaped lightweight engineering parts with exceptional mechanical properties. However, its remarkable properties, due to the rapid solidification and cooling rates during the LPBF process329microstructure of Ti6Al4V material have challenges to be addressed. Additionally, heat dissipation differences between the building direction and building plane make the Ti6Al4V microstructure more complicated. In the LPBF process when the laser heat source interacts with the metal powder melting and solidification occurred in a very confined region with a very rapid time manner.

In the LPBF process when the laser heat source interacts with the metal powder melting and solidification occurred in a very confined region with a very rapid time manner. After the laser scan, a substantial amount of heat dissipates through the material underneath the melt pool, resulting in a directional grain growth during the solidification. Thus, the microstructure of Ti6Al4V processed by LPBF technology has non-uniform grain structure. The microstructure variation is the fundamental source of the anisotropic mechanical behavior of the LPBFed Ti6Al4V material which is one of the most frequently studied properties of the material [5]. A great number of reported studies were seeking a solution to address this issue through the application of post-heat treatment (HT) [6] which requires additional research effort to optimize the effect on microstructure.

The application of HT increases the production cost with the necessity to advance equipment such as a hot isostatic pressure (HIP) furnace or controlled environment HT furnace. To avoid the additional cost of HT on LPBFed Ti6Al4V material, preheating laser scan, and post-heating laser scan studies were proposed to improve microstructure uniformity by controlling the heat input and cooling rates during the fabrication.In this study, the proposed innovative solutions of preheating laser scan and post-heating laser scan applications during the LPBF fabrication were evaluated. An additional laser scan was applied with a variety of laser scan speeds prior to the melting scan for preheating while a sequent laser scan was applied at a variety of scan speeds upon the melting scan to regulate the directional grain growth.

Results showed that both applications had a significant effect on the microstructure and mechanical response of the material and their effects were evaluated and compared in the presented study.2. Materials and MethodsThe material used in this study was gas-atomized Ti6Al4V Grade 5 powder, supplied by EOS North America (Pflugerville, TX, USA). The chemical composition(wt.%) of the powder consisted of 5.50 - 6.75 Al, 3.50 - 4.50 V, 0.20 O, 0.05 N, 0.08 C, 0.015 H, 0.30 Fe, and balance Ti, in accordance with ISO 5832- 3, ASTM F1472, ASTM F2924, and ASTM F3302 standards. The powder has a particle size distribution from 20 µm to 80 µm [7].The samples with a rectangular prism shape (Figure 1) were fabricated by LPBF technology on an EOS M290 SLM printer (EOS GmbH, Electro Optical Systems, Krailling, Germany) with a capacity of 400W Yb: YAG fiber laser.

The building plate was preheated to 80°C to reduce the thermal gradient at the initial layers of the fabrication.In this study, the effect of the innovative approaches of layerwise preheating and post-heating laser scan applications on the Ti6Al4V material’s microstructure and mechanical behavior was compared and discussed in detail. In both applications, the powder material was exposed to laser scan twice. Only the reference sample was exposed to a single laser track with the default process parameters of the LPBF process which were 280W laser power, 1300mm/s laser scan speed, 40 µm layer thickness, 67° scan rotation, 100 µm laser spot diameter and 100 µm hatch spacing.For preheating application, samples were exposed to a laser scan prior to the melting scan of default parameters whereas post-heating samples were exposed to a laser scan upon the melting laser scan.

For both applications, the effect of five different laser scan speeds at constant laser power with a total of five variations each was evaluated.330Figure 1. Schematic view and the orientation of the rectangular prism shape samples. The cartesian coordinates show the building direction along the Z-axis. The microstructural evaluation and mechanical testing were conducted along the ZX plane and XY plane to distinguish the effect of the orientation on microstructure and mechanical properties.Microstructure characterization was conducted to reveal the microstructure properties of porosity and α-phase lath thickness. A HITACHI S-3000N Variable Pressure Scanning Electron Microscopy (SEM) was used for microstructure imaging. Samples were cold-mounted before grinding and polishing.

A HITACHI S-3000N Variable Pressure Scanning Electron Microscopy (SEM) was used for microstructure imaging. Samples were cold-mounted before grinding and polishing. E-prep 4TM automated polisher was used for grinding (320 to 1200 Grit size silicon carbide abrasive discs) and polishing (1 µm diamond suspension on DiaMat polishing cloth and 0.04 µm colloidal silica for final polishing on a Chem-pol polishing cloth) with an individual force and rotation speed of the power head. The samples were cleaned with micro-organic soap and rinsed with isopropyl alcohol. The samples were etched before microstructure characterization by Kroll’s reagent (1-3 mL, HF, 2-6 mL HNO3, 100 mL water) to depict the lath structure and grain boundaries.An image analysis software Image J [8], was employed to quantize the porosity percentage and the thickness of the α-phase laths. The porosity level of the microstructure was evaluated in at least three regions of two magnification levels (500x and 800x).

The porosity level of the microstructure was evaluated in at least three regions of two magnification levels (500x and 800x). For lath thickness measurement microstructure images were first converted to RGB stack-type grayscale images prior to adjusting the contrast level. Then the threshold was adjusted by utilizing the auto-threshold comment of the software to create distinct contrast between individual laths. Particle analysis was employed between the range from 0.2 µm to 2.0 µm and the total particle counts were summarized for the mean value. This calculation was performed on at least six images at different magnification levels between 800x to 1500x for both XY and ZX planes.LECO LM 300 AT Micro Hardness tester equipment was used for the microhardness test, 500 g load was applied for 10 seconds in accordance with ASTM E384-22 [9].

Samples were polished including the final polishing step with 0.04 µm colloidal silica before the hardness test and at least 12 indentations were conducted for each sample.3. Results and Discussions3.1 Microstructural Analysis3.1.1 PorosityThe porosity levels of each sample were calculated by processing the SEM micrographs from at least three different regions that were captured at the magnification levels mentioned in Section 2. The porosity level of the reference sample without any preheating or post-heating laser scan was observed at 0.47% ± 0.015 and depicted with a black triangle marker in Figure 2.331A remarkable improvement in the LPBF Ti6Al4V material’s relative density was observed with the application of each preheating and post-heating laser scan. The authors observed the lowest porosity values at the defined laser power values of 252 W and 140 W for preheating and post-heating applications relatively.

The authors observed the lowest porosity values at the defined laser power values of 252 W and 140 W for preheating and post-heating applications relatively. The effect of the primary laser scan speed, during the preheating process, on porosity is shown in black (Fig. 2). All the preheating applications were conducted at a constant laser power of 252 W, and apart from the scan speed, all the process parameters were set to the default parameters mentioned in Section 2. The porosity results of the preheating laser scan speed of 650 mm/s, 975 mm/s, 1300 mm/s, 1625 mm/s, and 1950 mm/s were calculated.

The porosity results of the preheating laser scan speed of 650 mm/s, 975 mm/s, 1300 mm/s, 1625 mm/s, and 1950 mm/s were calculated. It was observed that the additional laser scan prior to the melting scan decreased the porosity level up to 0.01 %.2.42.4)g n i t l eM + )W 2 5 2 (n a c S g n i t a e h e r P : y t i s o r o P)%(21.61.20.80.421.61.20.80.4W 0 4 1 (n a c S g n i t a e Ht s o P + g n i t l eM: y t i s o r o P)%(005007009001500 Laser Scan Speed (mm/s)1100130017001900Figure 2.

The effect of the preheating laser scan speed on porosity at a constant preheating laser power of 252 W (in black) and the effect of the post-heating laser scan speed on the porosity at a constant post- heating laser power of 140 W (in red).A decremental trend in porosity level was observed in the case of preheating applications with higher preheating laser scan speeds at lower energy inputs resulting in a lower amount of porosity. An excessive porosity amount was observed at the preheating laser scan speed of 650 mm/s. The energy density of this preheating laser scan speed was 80.77 J/mm3 which was 1.8 times higher than the manufacturer’s optimized melting energy density (44.87 J/mm3). This redundant heat input formed an unstable melt pool with a higher amount of metal evaporation which increased the recoil pressure. Unbalanced recoil pressure provoked molten metal expulsion as spatters [10].

Unbalanced recoil pressure provoked molten metal expulsion as spatters [10]. A higher amount of spattering at the excessive energy input level contributed to a higher porosity level [11] in the microstructure of the preheated sample. Contrary to the lower scan speed rate, almost porosity-free microstructures (relative density ∼99.99%) were obtained at the higher preheating laser scan speeds of 1300 mm/s, 1625 mm/s, and 1950 mm/s.Similar to the preheating application, process parameters were kept identical to the default parameters except for the laser power and laser scan speed during the post-heating scan upon the melting laser scan. It was observed that the inherent porosity level of the microstructure decreased with the application of the post-heating laser scan for each scanning speed. The lowest porosity amount was observed at the post- heating scan speed of 650 mm/s with a percentage of 0.047% ± 0.038.

The lowest porosity amount was observed at the post- heating scan speed of 650 mm/s with a percentage of 0.047% ± 0.038. However, the improvement in porosity was limited in post-heating application compared to preheating, it was still 50% less than the initial microstructure of the reference sample.332The process-induced inherent porosities have a detrimental effect on the mechanical properties of LPBFed Ti6Al4V materials [12]. It is known that a complementary process of hot isostatic pressure (HIP) was applied to as-build LPBF Ti6Al4V to improve the porosity level of the microstructure [13]. In this study, a similar improvement in the porosity level delivered by HIP post-process was observed for both preheating-laser-scanned and post-heating-laser-scanned Ti6Al4V microstructure.

In this study, a similar improvement in the porosity level delivered by HIP post-process was observed for both preheating-laser-scanned and post-heating-laser-scanned Ti6Al4V microstructure. This can be considered as a promise by the application of additional laser scans as an alternate procedure for the expensive and time-consuming post-process of HIP to decrease the inherent porosity level of the LPBFed Ti6Al4V microstructure. Additionally, this promise of the proposed applications has the potential of superior mechanical properties of elongation [14] and fatigue strength [15] with lower process-induced defects.3.1.2 Grain StructureThe α + β titanium alloy Ti6Al4V microstructure was characterized and it was detected that the majority of the LPBFed material’s microstructure consisted of α-phase at room temperature. As the dominant phase, it is rational to evaluate the α-phase properties to correlate the microstructure and the mechanical response of the material.

As the dominant phase, it is rational to evaluate the α-phase properties to correlate the microstructure and the mechanical response of the material. α-phase has a basketweave-like grain structure in which the lath thickness contributes to the material's properties [16]. Figure 3 shows the lath structure of the samples along the XY and ZX planes for reference samples ((a) and (b)), preheating laser scan application samples ((c) and (d)), and post-heating laser scan application samples ((e) and (f)).(a)(b)(c)(d)333(e)(f)Figure 3.

SEM microstructure of the reference sample along the XY plane (a), along the ZX plane (b), Microstructure of the preheating laser scan application at 252W laser power and 1950 mm/s scan speed along the XY plane (c), and along the ZX plane (d), Microstructure of the post-heating laser scan application at 140W laser power and 1950 mm/s scan speed along the XY plane (e) and along the ZX plane (f).The variety in the microstructure of the reference Ti6Al4V between the XY and ZX planes was because of the directional solidification that occurred during the fabrication. Since a significant amount of heat is dissipated through the fabricated material underneath the melt pool, the cooling rate along the building direction (z-axis) was higher compared to the other axes. Thus, the microstructure along the different axes had different grain morphology.

Thus, the microstructure along the different axes had different grain morphology. In the presented study, this nonuniformity of the microstructure along the different orientations was investigated by analyzing the constituent α-phase lath structures. Figure 4 shows the effect of the preheating and post-heating laser scan on α-phase lath thickness.

(a) The effect of the preheating laser scans on the α-phase lath thickness variation between the XY and ZX planes. (b) The effect of the post-heating laser scans on the α-phase lath thickness variation between the XY and ZX planes.The non-uniform microstructure of the LPBFed Ti6Al4V is the fundamental reason for the anisotropic mechanical response of this material. The anisotropic mechanical properties are still a challenge that needs334to be addressed to utilize this technology further in industrial engineering applications. Previous studies [12,17] reported modifications in α-phase lath structure by employing post-heat treatments which required additional technical effort and cost to optimize the heat treatment process. The techniques in this study contributed to lath structure modification which resulted in lesser variation among the different orientations.

The techniques in this study contributed to lath structure modification which resulted in lesser variation among the different orientations. This was considered as the proven result of the preheating and post-heating laser scan applications on the improved isotropy of the LPBFed Ti6Al4V microstructure and mechanical properties.3.2 Mechanical TestingMicrohardness tests were conducted to analyze the effect of preheating and post-heating laser scans on the mechanical behavior of LPBFed Ti6Al4V. The reference sample’s hardness was measured as 406.3 ± 4.6 HV and 388 ± 4.4 HV along the XY and ZX planes respectively. Figures 5 (a) and (b) show the measured hardness values of the preheating and post-heating samples at different laser scan speeds.

(a) The effect of the laser scan speed on the microhardness of the LPBFed Ti6Al4V during preheating and post-heating laser scan along the XY plane. (b) The effect of the laser scan speed on the microhardness of the LPBFed Ti6Al4V during preheating and post-heating laser scan along the ZX plane. (c) Hardness difference between XY and ZX planes after the application of preheating laser scan speed at different scan speeds. (d) Hardness difference between XY and ZX planes after the application of post- heating laser scan speed at different scan speeds.It was observed that preheating laser scan speed has a significant effect on the hardness along the XY plane. The highest degradation in hardness was observed at the preheating laser scan of 1950 mm/s. The microstructure response after the preheating laser scan to the hardness along the ZX plane was not similar to the response along the XY plane.

The microstructure response after the preheating laser scan to the hardness along the ZX plane was not similar to the response along the XY plane. Although hardness decreased at the preheating laser scan speed of 1300 mm/s and 1625 mm/s, vice versa was found at the other scan speeds. The highest hardness improvement along the ZX plane with the preheating laser scan application was observed at the scan speed of 650 mm/s and the lowest degradation was observed at the scan speed of 1300 mm/s.Increasing post-heating laser scan speed from 975 mm/s to 1950 mm/s, the hardness of the LPBFed Ti6Al4V microstructure was increased compared to the reference condition along the XY plane. The highest degradation of hardness was observed at the scan speed of 975 mm/s. The hardness change after applying the post-heating laser scan along the ZX plane was not similar to the change on the XY plane.

The hardness change after applying the post-heating laser scan along the ZX plane was not similar to the change on the XY plane. Contrary to the response along the XY plane slight changes (<1%) in hardness were observed along the ZX plane after the application of the post-heating laser scan.Figures 5 (c) and (d) show that the application of preheating and post-heating laser scans decreased the hardness variation of the LPBFed Ti6Al4V microstructure between the XY and ZX planes at the reference condition. This observation was in parallel with the results discovered during the microstructure evaluation of the α-phase lath structures.In this study, it was observed that the challenge of the anisotropic mechanical response due to the non- uniform microstructure of LPBFed Ti6Al4V can be addressed with additional laser scans prior to, or upon the melting scan of the process.

This promise of the compared applications has the potential to skip the heat treatments applied after the fabrication. The effect of the compared applications on other microstructure features and mechanical properties of the LPBFed Ti6Al4V will be investigated by the authors in future studies.3364. ConclusionIn this study, the effect of the preheating laser scan and post-heating laser scan on LPBFed Ti6Al4V material’s microstructure was observed. The mechanical response also have been evaluated through hardness measurements and it has been concluded that post-heat treatment processes need not be required if they are used during fabrication. The findings of this study listed as: • Preheating laser scan application prior to the melting scan with the default parameters had a significant effect on the inherent porosity level of the LPBFed Ti6Al4V material.

It was observed that preheating scan application decreased the porosity level of the reference condition from 0.47% to 0.01%.Application of the preheating scan modified the microstructure by decreasing the α-phase lath thickness variation between XY and ZX planes and created a more uniform grain structure along different orientation. It was observed that the variation decreased to 1.3% between the planes. • Preheating laser scan application has a significant effect on the hardness values of the material. It was observed that the hardness variation between the XY and ZX planes decreased by less than 1.5%.Post-heating laser scan application upon the melting scan had a similar effect on the process- induced inherent porosity level as the preheating application.

Compared to the preheating, post- heating scan had a limited effect on porosity but it was still 50% less than the reference condition of the LPBFed Ti6Al4V material.Post-heating laser scan application improved the uniformity of the Ti6Al4V microstructure by decreasing the α-phase lath thickness variation between the XY and ZX planes. It was observed that the variation decreased to 3.1% between the planes which was not as significant as the preheating application.Post-heating laser scan decreased the anisotropic mechanical response by decreasing the variation of the hardness between the XY and ZX planes up to 3.7% for LPBF-processed Ti6AL4V material.5. References[1][2][3][4]S. Liu, Y.C. Shin, Additive manufacturing of Ti6Al4V alloy: A review, Mater Des. 164 (2019) 107552. https://doi.org/https://doi.org/10.1016/j.matdes.2018.107552. T.S. Tshephe, S.O. Akinwamide, E. Olevsky, P.A.

mparison of Material Properties and Microstructure of Specimens Built Using the 3D Systems Vanguard HS and Vanguard HiQ+HS SLS SystemsT. J. Silverman, A. Hall, B. South, W. Yong and J. H. Koo* Laboratory for Freeform Fabrication Department of Mechanical Engineering The University of Texas at Austin Austin, TX 78712 *Corresponding author: jkoo@mail.utexas.eduAbstractThe HiQ upgrade to the 3D Systems Vanguard selective laser sintering (SLS) machine in- corporates a revised thermal calibration system and new control software. The paper compares the tensile modulus, tensile strength, elongation at break, ﬂexural modulus, Izod impact resis- tance and microstructure of two batteries of standard specimens built from recycled Duraform PA (Nylon 12). The ﬁrst set is built on a Vanguard HS system and the second on the same system with the HiQ upgrade installed. The upgrade reduces user intervention, decreases to- tal build time and improves surface ﬁnish.

The upgrade reduces user intervention, decreases to- tal build time and improves surface ﬁnish. However, using the default processing parameters, tensile, ﬂexure and impact properties are all found to decline after the upgrade is installed.1IntroductionThe Vanguard HS selective laser sintering (SLS) machine can be upgraded to HiQ+HS by the manufacturer, 3D Systems, Incorporated. The upgrade comprises new infrared sensors for the feed bins, a calibrating blackbody mounted on the roller carriage and updated software that performs real-time temperature calibration and control. In previous work [1], we built a battery of mechani- cal test specimens from recycled DuraForm PA (Nylon 12) powder using a Vanguard HS machine, installed the HiQ upgrade, then repeated the build. The specimens were then tested with reference to the applicable ASTM standards. The microstructure of typical specimens was evaluated using scanning electron microscopy (SEM).

The specimens were then tested with reference to the applicable ASTM standards. The microstructure of typical specimens was evaluated using scanning electron microscopy (SEM). A summary of the previous work [1] is presented here, fol- lowed by interpretation and recommendations for industrial and research practices and for future work.Figure 1: The specimen battery used to quantify the upgrade performance.2 Experimental procedureThe specimen battery contains a total of 240 ASTM standard tensile, ﬂexure and impact test speci- mens [2-4], ten in each possible orientation (Figure 1). The specimens are arranged as close to the center of the build area as possible. They are uniquely numbered so that the effects of position and orientation on material properties can be analyzed.The battery was built using DuraForm PA (Nylon 12) material recycled from part cake from a local service bureau.

The powder was blended and its consistency checked using a prescribed procedure based on melt ﬂow index (MFI) [2].The specimen batteries were built using the 3D Systems recommended parameter values, shown in Table 1. The part bed temperature for the HS build was selected manually using the “ﬁnd temps” procedure. Because of its self-calibration capability, the HiQ+HS machine does not require the “ﬁnd temps” procedure to be performed for materials with a built-in conﬁguration ﬁle, including DuraForm PA. For materials without predeﬁned parameter conﬁgurations, the procedure must only be performed once.Table 1: Manufacturer-recommended processing parameters were used for both specimen builds.Part Heater Set Point Left and Right Feed Heater Set Point Fill Laser Power Powder Layer ThicknessHS HiQ+HS 152.0 C 173.0 C 130.0 C 140.0 C 42 W 0.004 in 0.004 in42 WThe HS build lasted 26 hours, while the HiQ+HS build lasted 20 hours.

The HS build had a notably stiffer part cake that was harder to remove than the ﬂour-like part cake in the HiQ+HS build. The surface ﬁnish of the pre-upgrade specimens was noticeably rougher than that of the post-(cid:61)1174,Σ(cid:61)1051014151910(cid:61)1558,Σ(cid:61)66post(cid:45)upgradexx136148887167741045x14121519661203481082101213611519pre(cid:45)upgrade10451361(cid:61)1499,Σ(cid:61)85104588761519x12216771045614x1412031414167714120313612108871045146(cid:61)1595,Σ(cid:61)5061248120316771361167712031214(cid:61)1123,Σ(cid:61)151xorientationyorientationzorientation141412148104528874288721519(cid:61)1473,Σ(cid:61)4016771203410136188715198xFigure 2: Histograms for tensile modulus (MPa) by orientation for both builds.upgrade specimens.

Surface powder was removed from the specimens using glass bead abrasive. The specimens were then tested with reference to the appropriate ASTM test procedures [2-4].3 Results and discussionThe previously published values [1] for tensile modulus and elongation at break are incorrect due to an equipment limitation which led to a deviation from the ASTM tensile test procedure [3]. The values given here have been corrected and can be compared to each other. Illustrative comparisons can be made to other data collected using the same standard, but detailed quantitative comparison may not be valid. The material data sheet to which the previous work [1] refers has been shown to contain errors [4].

The material data sheet to which the previous work [1] refers has been shown to contain errors [4]. The comparisons here are to the old data sheet [5], the corrected data sheet [6], which may contain a spurious impact resistance ﬁgure, and the ASTM speciﬁcation for injected or extruded general purpose Nylon 12 [7].3.1 Tensile propertiesTensile properties show strong variation with respect to orientation in the build volume. His- tograms for tensile modulus (Figure 2), ultimate tensile strength (Figure 3) and elongation at break (Figure 4) are separated by orientation.The tensile properties are best in specimens with the y orientation, that is, with their longest dimension along the build volume’s y axis.

Elongation at break is shown to be below the speciﬁed value both before and after the upgrade.All three tensile properties that were measured substantially declined between the two builds studied here. This indicates that, with the default, manufacturer-recommended processing param- eters, the upgrade may not result in improved tensile properties.3.2 Flexural modulusFlexural modulus does not vary substantially with the orientation of the test specimen in the build volume. Histograms for ﬂexural modulus are shown in Figure 5.Table 2: Comparison of overall mean of measured material properties with manufacturer and ASTM speci- ﬁcations.Tensile modulus (MPa) Ultimate tensile strength (MPa) Elongation at break (%) Flexural modulus (MPa) Impact resistance (J/m)HS 1550 48 7 1860 242Old 3D HiQ+HS spec. [5] 1256 37 5 1383 2111600 44 9 1285 214New 3D ASTM spec. [6] 1586 43 14 1387 32spec.

Again, with the recommended processing parameters, the upgrade alone is not adequate to produce an improvement in ﬂexural modulus.3.3Impact resistancePre-upgrade impact resistance is not sensitive to orientation, but specimens built in the z orientation after the upgrade have relatively low impact resistance. Histograms for impact resistance are shown in Figure 6.The overall mean impact resistance ﬁgures in Table 2 conﬁrm that impact resistance also declined between the two builds. The upgrade does not improve impact resistance when the manufacturer-recommended parameters are used.3.4 DensityThe specimens’ approximate densities were measured by hand using water displacement and a balance. These densities are plotted according to x-y position in the build volume in Figure 7.There is no clear trend in density with position.

These densities are plotted according to x-y position in the build volume in Figure 7.There is no clear trend in density with position. Density declines slightly between the twobuilds, but this effect may not be signiﬁcant compared to the measurement error.10.851.05113030.90××××××x-coordinate (in)y-coordinate (in)20××02132231.1511.123density (g/cc)31××1230.952130.81××x-coordinate (in)y-coordinate (in)2××1××Figure 7: Density map (top view) for the pre-upgrade (left) and the post-upgrade (right) builds.

The density for all specimens (n = 240) was 0.997 ± 0.048 g/cc (pre-upgrade) and 0.970 ± 0.056 g/cc (post-upgrade) ( ¯x ± σ ).Figure 8: Scanning electron micrographs of new DuraForm PA powder (left) and the recycled part cake used for this work (right).3.5 MicrostructureA sample of the recycled part cake from which the test builds were fabricated was imaged using scanning electron microscopy (SEM) and compared to a micrograph of fresh material. The recy- cled and new powder have visibly similar texture, particle size and size distribution, as shown in Figure 8.SEM micrographs of tensile-test fracture surfaces of pre- and post-upgrade tensile specimens were used to characterize their microstructure (Figure 9). The surfaces show visibly similar occur- rence of voids and cracking.

The surfaces show visibly similar occur- rence of voids and cracking. Micrographs of cryogenic fracture surfaces, created by breaking the specimens after chilling them with liquid nitrogen, are also visibly similar (Figure 10).Substantial differences in microstructure between new and recycled SLS powder and betweenpre-upgrade and post-upgrade fracture surfaces are not apparent.398Figure 9: Micrograph of the fracture surfaces for tensile specimens built before (left) and after (right) the upgrade.Figure 10: Micrograph of cryogenic fracture surfaces of tensile specimens built before (left) and after (right) the upgrade at 500× (top) and 1000× (bottom) magniﬁcation.4 Summary and conclusionWhen manufacturer-recommended processing parameters were used, all of the mechanical prop- erties measured in this study declined upon the installation of the upgrade.

However, these recom- mended parameters are intended to be a starting point and not a formula for optimal mechanical properties. They are chosen by the manufacturer to match the ASTM speciﬁcations for injection- molded Nylon 12, as a compromise between better material properties and easier breakout [4]. Indeed, both the pre-upgrade and post-upgrade builds exceed the ASTM speciﬁcations for general purpose Nylon 12 in those properties that are speciﬁed [7]. No substantial changes in microstruc- ture were observed before and after the upgrade, indicating that no qualitative microstructural dif- ference accounts for the decline in mechanical properties. This suggests that mechanical properties could be improved by adjusting machine settings away from the manufacturer’s initial recommen- dations.Each SLS machine, even of a particular model, performs differently [4]. Specifying the pro- cessing parameters that result in the optimal mechanical properties is currently an iterative process.

Specifying the pro- cessing parameters that result in the optimal mechanical properties is currently an iterative process. Like the operation of any complex machine, this optimization process involves a great deal of oper- ator knowledge and experience. The HiQ upgrade does not eliminate this process [4]. Parameters that are optimal for a HS system are not optimal for a HiQ+HS system.Dozens of SLS processing parameters are available for adjustment (Table 3) and many are subject to measurement error. The upgrade reduces error in the measurement of temperatures by providing real-time sensor calibration. This eliminates the ﬁnd temps procedure for experimentally ﬁnding the fusion temperature of the part bed and the temperature ramping procedure used to compensate for sensor drift during the build.

This eliminates the ﬁnd temps procedure for experimentally ﬁnding the fusion temperature of the part bed and the temperature ramping procedure used to compensate for sensor drift during the build. This makes builds faster to set up and run and is expected to support consistency between builds.5 Recommendations and future workFinding the best parameters for a particular SLS machine is still an open-ended, iterative process, even after the upgrade is installed. We recommend that the institutions researching, developing, manufacturing and using SLS machines cooperate to develop and publish a procedure for searching the range of processing parameters and optimizing mechanical properties.

The procedure, or a prototype thereof, should be used on a pre-upgrade machine and a batch of test specimens should be manufactured from fresh, unrecycled material using the resulting optimized parameters. The same optimization and specimen build should then be performed on an upgraded machine. This will fairly quantify any beneﬁts of the upgrade and demonstrate the feasibility and efﬁcacy of the optimization procedure.Further testing and the development of at least part of the optimization procedure is planned atthis laboratory.AcknowledgementsThe authors would like to thank Kevin McAlea, Simon van de Crommert, Kevin Marshall, Art Delgado, Scott Danniel and Jennifer Torkelson of 3D Systems Corporation for providing a HiQ upgrade and subsequent support for the Vanguard HS SLS machine to the Laboratory for Freeform Fabrication. Their continuous support of our SLS research activities at the University of Texas at Austin is greatly appreciated.

Their continuous support of our SLS research activities at the University of Texas at Austin is greatly appreciated. We would also like to thank Kent Firestone of Solid Concepts for advising us on the SLS build conﬁguration as well as his helpful guidance in our SLS processing.References[1] T.J. Silverman, A. Hall, B. South, W. Yong, and J.H. Koo. Material properties and microstruc- tural characterization of specimens manufactured by selective laser sintering. SME Technical Paper TP07PUB58, Society of Manufacturing Engineers, Dearborn, Mich., 2007.[2] T.J. Gornet, K.R. Davis, T.L. Starr, and K.M. Mulloy. Characterization of selective laser sintering materials to determine process stability. In Solid Freeform Fabrication Proceedings, pages 546–553, 2002.[3] ASTM International. ASTM D 638-03: Standard Test Methods for Tensile Properties of Plas-tics.[4] Scott F. Danniel and Jennifer Torkelson of 3D Systems Corp., Rock Hill, S.C.

ASTM D 638-03: Standard Test Methods for Tensile Properties of Plas-tics.[4] Scott F. Danniel and Jennifer Torkelson of 3D Systems Corp., Rock Hill, S.C. Personal Com-munication, May 2007.[5] 3D Systems Corporation, Rock Hill, South Carolina (formerly in Valencia, California). Du-raform PA and GF Plastic, March 2005.[6] 3D Systems Corporation, Rock Hill, South Carolina (formerly in Valencia, California). Du-raform PA Plastic, November 2006.[7] ASTM International. ASTM D 4066a: Standard Classiﬁcation System for Nylon Injection andExtrusion Materials (PA)..

Microstructures and microstructural issues are examined by optical metallography, SEM, TEM, EDS, and XRD while mechanical properties included selective specimen tensile testing and Vickers microindentation (HV) and Rockwell C-scale (HRC) hardness measurements. Orthogonal (X-Y) melt scanning of the electron beam during AM produced unique, orthogonal and related Cr23C6 carbide (precipitate) cellular arrays with dimensions of ~2μm in the build plane perpendicular to the build direction, while connected carbide columns were formed in the vertical plane, parallel to the build direction.INTRODUCTIONThe evolution of manufacturing technologies implies the need of characterizing different materials available in order to obtain the best properties intended for the specific applications. Rapid manufacturing by electron beam melting is becoming an ideal technology to create complex shapes through computer-controlled self-assembly by sintering or melting of powder layers [1].

Rapid manufacturing by electron beam melting is becoming an ideal technology to create complex shapes through computer-controlled self-assembly by sintering or melting of powder layers [1]. Since traditional manufacturing consists of creating a final shape by the use of substractive or formative processes, rapid prototyping consists of forming a model, one layer at a time, from bottom to top [2].

By combining 3D implant models and rapid prototyping technologies gives us the capability of fabricating custom anatomical implants [3].Nowadays some manufacturing enterprises have started to use rapid prototyping methods (or additive manufacturing, AM) for complex pattern making and component prototyping to shorten the time for pattern, molds and prototype development [4].The powder material utilized for this project is Co-26Cr-6Mo-0.2C since it is an alloy that can withstand high temperatures, and has a high wear resistance, it can be used in gas turbines; valve seats, nuclear power plants, automobile engines, aerospace fuel nozzles and engine vanes and components, as well as in a variety of orthopaedic and dental implants [5-8].METHODS AND PROCEDURESElectron beam melting was utilized to fabricate CoCrMo specimens to be analyzed by optical microscopy and transmission electron microscopy, in addition, hardness and tensile testing was performed.

A schematic of the EBM system is shown in figure 1a, the EBM system is computer driven and works in vacuum. Number 1 shows the location of the electron gun operating at 60kV, the beam depicted in number 2 is focused and bent by electromagnetic scan coils. Number 3 shows the location of the powder to be used for the process while number 4 shows the building table that moves down ~100 m to provide a new layer of powder. The basic operation procedure of the EBM consists of preheating and then melting the desired areas in each layer, at an approximate temperature of 830 C, as directed by the STL file created. Once the components were fabricated they were optically analyzed in the vertical and the horizontal planes which can also be more specifically described as parallel and perpendicular to the building direction respectively. In addition to analyzing as- fabricated components an annealed and rough polished component in the shape of a femoral-knee implant was also analyzed.

In addition to analyzing as- fabricated components an annealed and rough polished component in the shape of a femoral-knee implant was also analyzed. The annealing heat treatment consisted of initial hot isostatic pressing (HIP) at ~1200 C for 4h in Ar at 103 bar, followed by quench from a homogenizing treatment at 1220 C for 4h in Ar, at 75 C/min. The homogenizing temperature was ~0.8 TM (~1430 C), and everything was performed following ASTM F75 CoCr Alloy standard [9].More than one etching solution had to be used to complete this study, to begin with a solution consisting of 6:1 HCL:H2O2 (3%) for 16 h for the as-fabricated components while the annealed component used a solution consisting of 6:1 HCL:H2O2 for an average time of 5 minutes.

The solution utilized for the TEM specimens consisted of 15% HClO4 and 85% acetic acid deposited in a jet polisher system (Tenupol-5) at a temperature range of 25-40 C and 20V.XRD spectra were obtained from a Brucker AXS-D8 system using a Cu target. A Vickers hardness (HV) indenter (25-100 gf (0.25-1N) load at ~10s load time in a Shimadzu HMV-2000 system and a Rockwell C-scale hardness (HRC) tester (1.5 kN load) were used for this project while tensile testing was performed in an Instron system at ambient temperatures (~ 20°C) at a strain rate of 3x10-3s-1.RESULTS AND DISCUSSIONFigure 1b shows CoCrMo powder used for this project, having a nominal powder size diameter of 40 m. A higher magnification SEM micrograph of the CoCrMo powder is shown in figure 2. Microstructural characterization was performed in the EBM fabricated components obtaining a unique and complex array of carbides as the one illustrated in figure 3.

Microstructural characterization was performed in the EBM fabricated components obtaining a unique and complex array of carbides as the one illustrated in figure 3. Figure 3 was obtained from the vertical plane of a cylindrical component (the arrow shows the building direction) with dimensions consisting of a diameter of 1.5 cm and ~11 cm in length. The vertical plane of any component stands for the parallel plane related to the building direction, while the horizontal plane stands for the plane perpendicular to the same. It is also important to mention that cylinders of these dimensions were utilized to machine tensile test specimens for this project.Figure 1. (a) EBM system schematic.

It is also important to mention that cylinders of these dimensions were utilized to machine tensile test specimens for this project.Figure 1. (a) EBM system schematic. (b) SEM image of CoCrMo powder particlesXRD was performed on the CoCrMo powder before being utilized by the EBM system, on the horizontal plane of as-fabricated cylindrical specimens and also on an annealed and rough polished section cut from a femoral knee implant component shape as illustrated in figure 4. It can be observed how the cylindrical component shows a variety of crystallographic and compositional phase mixtures, mostly hcp with a Co or CoCr fcc matrix. The carbide found is Cr23C6 which is an fcc with a lattice parameter of 10.66A°. It can be appreciated how the carbide peak is absent in the annealed and polished knee component, which is also observed from the metallographic images obtained for this specimen when comparing figure 5 to figure 6.

In figure 5 the regular carbide arrays representing precipitation are created by the cross-scanning of the electron beam during preheating and in suceessive layer building, with dimensions of ~2μm are observed in the horizontal plane while similarly spaced columns of carbides are shown extending along the build direction. Other regular and irregular columnar carbide features are observed in the vertical plane view in figure 5a with dimensions similar to those shown in the horizontal plane. Figure 6 shows the microstructure obtained from a small section of the annealed and polished component revealing an equiaxed fcc grain structure and annealing twins.Figure 2. High magnification SEM image of CoCrMo powderFigure 3. Vertical plane of a cylindrical component fabricated by EBM showing a unique and complex array of carbidesPOWDERaCYLINDERbPOLISHED KNEECOMPONENTcFigure 4.

It can be observed how Cr23C6 precipitates are present along with stacking faults and dislocations. From the selected area electron diffraction (SAED) pattern it can be observed the presence of the fcc Co matrix spots combined with the precipitate related spots belonging to the fcc (100) Cr23C6 pointed by the small arrow close to the beam stopper. Figure 8 shows a low and high magnification TEM image representing the vertical and horizontal planes of as-fabricated EBM components. A columnar array of Cr23C6 precipitates shows to be spaced approximately 100 to 200nm. It can also be appreciated from figure 8b that dislocation arrangements are intermingled with carbide precipitates as well as contrast fringes indicating linear stacking-fault features, present in both horizontal and vertical plane views. Figure 9 is a TEM image from an annealed and polished femoral knee component showing a greater stacking fault density in contrast to figures 7 and 8.

Figure 9 is a TEM image from an annealed and polished femoral knee component showing a greater stacking fault density in contrast to figures 7 and 8. It can also be observed from the SAED pattern that there is no evidence of Cr23C6 carbide precipitates and the (100) fcc SAED pattern exhibits a calculated lattice parameter of 3.55 Å, consistent with fcc Co.Figure 7. TEM bright-field image of the horizontal plane for a cylindrical component showing Cr23C6 precipitates, dislocations and stacking faults. The SAED pattern insert shows fcc Co (matrix) diffraction spots and (100) fcc Cr23C6 diffraction spots (arrow)Horizontal PlaneHorizontal PlaneVertical Planea)Vertical PlaneFigure 8. Shows a 3D representation of TEM images showing columnar arrays of Cr23C6 precipitates (a) and at a higher magnification (b)Figure 9.

**Rockwell C-scale hardness *vHorizontal plane (see figure 5) vAverage for 2 tests vvHorizontal plane hardness/Vertical plane hardness (see figure 5) v*v0.2% engineering offset yield stress.5.946--4.740----------Table 1 shows the tensile test results of specimens prepared from as-fabricated cylindrical components as well as the HV hardness test results for precursor powder, as- fabricated and annealed components as well as HRC hardness in the vertical and horizontal plane of cylindrical and block shape components and the annealed femoral- knee component. The yield stress is consistent with wrought and cast products, while UTS is considerably higher than as-cast or wrought ASTM F75 Co-Cr-Mo alloy, where nominally the yield stress and UTS are 0.5 GPa and 0.9 GPa respectively.

An average elongation of 3.6% was obtained from two tensile samples where one value was 1.9% and the other 5.3%, when compared to as-cast ASTM F75 CoCrMo the elongation is <1% while wrought CoCrMo has an elongation of ~5%.CONCLUSIONSThe EBM fabrication of components and prototypes from Co-26Cr-6Mo-0.2C powder (Co0.8Cr0.2 hcp crystal structure) creates Co-26Cr-6Mo-0.2C monoliths having fcc CoCr matrix with CrMo phase components and a unique, electron beam scan- produced Cr23C6 fcc orthogonal carbide array when viewed perpendicular to the build direction, and carbide columns connected to these arrays when viewed in a plane parallel to the build direction. In the same manner, after annealing, an equiaxed, fcc CoCr grain structure containing {111} coincident annealing twin forms with Cr23C6 carbides mainly in high energy grain boundary positions. TEM images demonstrate a high density of intrinsic stacking faults on {111} planes, and no matrix carbides.

TEM images demonstrate a high density of intrinsic stacking faults on {111} planes, and no matrix carbides. Tensile testing of as- fabricated EBM cylindrical components showed improved properties compared to wrought or cast Co-26Cr-0.6Mo alloys such as ASTM F75.REFERENCES[1] Murr, L. E., et al. 2009. Microstructure and mechanical behavior of Ti-6Al-4V produced by rapid-layer manufacturing, for biomedical applications. Journal of the Mechanical Behavior of Biomedical Materials, Vol 2, (1). 20-32.[2] Kamrani, A. K., Nasr E. A., 2006. Rapid Prototyping. Theory and Practice. Springer Science+ Business Media, Inc.[3] Vail, N. K., et al. 1999. Materials for biomedical applications. Materials and Design 20. 123-132.[4] Yan, X. and Gu, P. 1996. Survey: A review of rapid prototyping technologies and systems. Computer-Aided Design. Vol 28, No. 4. 307-318.[5] Antony, K. C. 1983. Wear-resistant cobalt base alloys. J. metals 35 (1983) 52-60.[6] Alamert, S., and Bhadeshia, H.K.D.H.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA Comparison of Modeling Methods for Predicting the Elastic-Plastic Response of Additively Manufactured Honeycomb StructuresRaghav Sharma1, Thao Le1, Jiaxu Song1, Ethaniel Harms1, Daniel Sowa1, Alex Grishin2, Dhruv Bhate1 1Arizona State University, Mesa, AZ 85212 2Phoenix Analysis & Design Technologies, Inc., Tempe, AZ 85284AbstractValid and accurate models describing the mechanical behavior of additively manufactured cellular materials are crucial to enabling their implementation in critical-to-function parts. Broadly speaking, the modeling approaches commonly used in the literature fall into three categories.

Broadly speaking, the modeling approaches commonly used in the literature fall into three categories. Each of these differs in the level of discretization at which the cellular behavior is modeled: at the level of each material point, at the level of the unit cell or at the level of a connecting member that constitutes a unit cell. Each of these three approaches relies on different characterization techniques and the way in which the resulting data is leveraged in the development of the model. In this work, we critically examine all three modeling approaches using FEA and compare their accuracy in the prediction of the elastic and plastic behavior of experimentally characterized hexagonal honeycomb structures made with Fused Deposition Modeling, and discuss the pros and cons of each method.IntroductionModeling, for the purposes of this paper, refers to the methods of prediction of the behavior of cellular materials.

Modeling is important for at least two reasons: first, it enables the translation of predictability from a lower level of complexity (like a tensile test specimen) to the application at hand (such as a rotating fan blade). This is a key step in validation since not everything can be tested in every environment possible. Modeling also serves another purpose - it drives optimization and allows us to improve performance. Models allow us to gain insight into the design variables that matter and enable the evaluation of multiple design options in search of the optimal one.For a model to be of practical value, it needs two key ingredients working together. The first of these is a mathematical description that relates behavior to design and material parameters (this could be analytical and/or leverage computational tools such as Finite Element Analysis - FEA). Secondly, the model needs a quantified material representation that can be implemented into the mathematical description.

Secondly, the model needs a quantified material representation that can be implemented into the mathematical description. In this paper, we focus on the different ways cellular materials can be represented for use in FEA software – in other words, we investigate how the material properties can be determined and represented for accurate results.A Review of ApproachesMaterial modeling approaches for cellular materials fall into three different categories, depending on the level of discretization at which the property is modeled. Arranging these in order of the scale at which material behavior is represented, these models can operate at the level of each material point (bulk property models), at the level of the connecting member (member models) or791finally, at the level of the cell (homogenization models). In the following sections, we deal with each of the three in turn, but in a slightly different order to enable meaningful comparisons.1.

In the following sections, we deal with each of the three in turn, but in a slightly different order to enable meaningful comparisons.1. Bulk PropertyThe most straightforward approach in modeling cellular materials is to use bulk material properties to represent what is happening to the material at the cellular level. This approach does away with the need for any cellular level characterization and in so doing, does not have to account for size effects or other experimental artifacts that are specific to cellular materials. ASTM or ISO standards can be leveraged to develop databases of material behavior and this allows for a proper statistical basis to be established. However, the assumption that the connecting struts/walls in a cellular structure behave the same way the bulk material does can particularly be erroneous for AM processes that can introduce significant size-specific behavior and large anisotropy.

There are two reasons for this – the first has to do with the process dependence inherent to most AM processes, the second has to do with the geometry related variations specific to cellular materials themselves.Most AM processes have a strong orientation dependence to their mechanical properties, with the build direction being typically the weaker, in comparison to the in-plane properties. These differences are strongly process dependent, with parts made with Fused Deposition Modeling (FDM) processes being strongly anisotropic on one end, and Electron Beam Melting parts being less so at the other end of the spectrum. However, anisotropy can be represented in orthotropic material models to mitigate the impact of these differences. What is harder to represent is the dependence of the properties on the meso- and/or microstructure, which in turn often has strong dependence on the process settings used to realize that part.

What is harder to represent is the dependence of the properties on the meso- and/or microstructure, which in turn often has strong dependence on the process settings used to realize that part. In the example shown in Figure 1, for the FDM process, four ASTM D638 standard specimens were manufactured with the identical geometry, but varying process parameters – in this case, the number of contours and rasters in the part was varied. The resulting load-displacement curves are significantly different to each other, implying process dependent elastic moduli and failure stresses.Figure 1. Effect of processing parameters (number of contours) on tensile load-displacement for the Fused Deposition Modeling processThe reason process dependence matters particularly for cellular materials comes down to their typically fine features. For a thin strut or wall, the additive process may only pass through a792given feature a few times in each layer.

For a thin strut or wall, the additive process may only pass through a792given feature a few times in each layer. In many cases, 1 or 2 contours are all that is needed to create the feature at that layer. As shown in Figure 2 for thin walls made with laser powder bed fusion (LPBF), as the thickness of the feature increases, additional passes are needed, which in turn can change the microstructure and mechanical behavior of the cellular material as an aggregate.Figure 2. As feature thickness increases, more passes of the laser are needed, resulting in modifications to structure (shown here for the Laser Powder Bed Fusion process)Most additively manufactured cellular materials tend to have feature dimensions on the order of tens to hundreds of microns. Most of the processes are themselves capable of printing parts on the order of hundreds of millimeters.

Most of the processes are themselves capable of printing parts on the order of hundreds of millimeters. This combination of resolving fine features in relatively large parts is one of the attractive aspects of AM technologies, but does introduce a challenge from a modeling perspective. Consider the honeycomb structure in Figure 3, for example, manufactured on a production grade FDM system (the Stratasys Fortus 400mc) [1]. Measurements of the wall thicknesses show an average error between measured and designed thicknesses of 100 µm, which is within the machine specifications and ordinarily is not noteworthy. However, this minor variation amounts to a 7% error in the true thickness of the honeycomb walls as designed. In honeycombs, which are bending-dominated for thin wall structures, the effective modulus of the structure scales by the cube of the thickness, so this would be a very large source of error in modeling the honeycomb.

For stretch-dominated structures such as lattices, modulus scales linearly, so this would still amount to a significant error. Therefore, dimensional errors that are within specifications at the level of bulk structures including typical ASTM/ISO test specimens, are significant first-order errors when the dimensions get as small as they do. It is for this reason that models that exclude these effects, as bulk property models do, are unlikely to be good predictors of behavior, unless corrections are made to the model to account for them.Dimensions are only one of the sources of imperfection that bulk property models do not take into consideration. The others include surface roughness, waviness, defects like pores and cross-section variations by orientation, all of which are discussed for the LPBF process in a paper by Liu et al. [2].

The others include surface roughness, waviness, defects like pores and cross-section variations by orientation, all of which are discussed for the LPBF process in a paper by Liu et al. [2]. Some researchers use bulk property assumptions but then develop correction793factors to get good agreement between model and experiment, as demonstrated for Selective Laser Sintering (SLS) process by Neff et al. [3] and the LPBF process by Smith et al. [4].(a)(b)(c) Figure 3. (a) A honeycomb sample, (b) optical scan image using blue light scanning and (c) 12-sample data showing a mean of 1.625mm against a designed value of 1.524mm – a 7% error in thickness, adapted from Le et al. [1]2. HomogenizationHomogenization involves estimation of an effective property of the cellular structure without explicit regard to the cellular geometry itself.

[1]2. HomogenizationHomogenization involves estimation of an effective property of the cellular structure without explicit regard to the cellular geometry itself. These properties are determined experimentally through characterization of cellular materials, and in that sense represent an improvement over the bulk property method. Another advantage inherent to homogenization is that this approach has significantly lower computational expense associated with its implementation in simulation software since the cellular structure is not explicitly modeled in the computational environment but instead replaced with a solid structure that has effective properties, as shown in Figure 4.

Additionally, it is relatively straightforward to develop a model by fitting a power law to experimental data as discussed before and shown in the equation below, relating the effective modulus E* to the bulk material property Es and their respective densities (ρ and ρs), by solving for the constants C and n. Similar equations can be developed and fitted to the observed data for other material parameters of interest.(1)𝐸𝐸∗= 𝐶𝐶 𝐸𝐸𝑠𝑠 �𝜌𝜌 𝜌𝜌𝑠𝑠 �𝑛𝑛794Figure 4. (left) Homogenization enables the replacement of a cellular material with a solid of effective properties, (right) which can greatly reduce computational expense when simulating engineering structuresThe homogenization approach has been used by several researchers in the literature [5]– [8] and while it has the advantages mentioned above, it also has some difficulties in being used as a reliable material model in analysis & simulation that need to be taken into consideration.

An immediate limitation of homogenization is that it is dependent on the cell shape – each shape needs to be individually characterized and relationships developed. Further, this needs to be done at different volume fractions, otherwise the model is limited to one specific size for a given shape. This increases the experimental effort required significantly, in comparison with the other approaches discussed in this section. A second, more serious limitation arises from the fact that for a given shape, there are two ways to modulate the volume fraction – change the size of the cell, or change the thickness of the struts or walls that make up the cell. It has been shown that these two ways of modulating the volume fraction do not result in similar calculations of modulus and other properties – in other words, homogenization models based on relative density alone are prone to error.

Finally, it is not clear if homogenization models can predict well the behavior of structures with non-uniform cell size since transitions between different relative densities can introduce shapes that do not conform to the shape assumptions in the model. Nonlinear behavior, damage and fracture are also harder to model since these are so geometry dependent, transient phenomena.All these limitations when taken together, place serious constraints on homogenization models for cellular materials. While they are likely to serve well when the shape and size of the cellular material is fixed and known, they do not lend themselves well to the spirit of AM and the design freedom it enables.3. Member ModelingThe two above approaches seek to represent cellular material behavior at two extreme ends: the bulk property approach seeks to establish a point-wise material representation whereas the homogenization approach seeks to do so at the level of the unit cell.

The third approach, which we may term member modeling, involves describing behavior at a level in-between, viz. the connecting member or element that networks to constitute the cellular structure. This is done by identifying an elemental member that constitutes the cellular material and designing and testing a specimen that embodies that member, as shown in Figure 5.795Figure 5. Member modeling involves identifying the key member contributing to the deformation and characterizing itThe basic notion of the member modeling method (though with different terminology such as micro-strut for lattice materials) has been used by several researchers with mixed results [1], [9]–[13]. The advantage of this approach is that if we are able to accurately describe behavior at the level of a connecting member, it enables a truly shape independent model [1].

The advantage of this approach is that if we are able to accurately describe behavior at the level of a connecting member, it enables a truly shape independent model [1]. Additionally, more insight can potentially be gained by studying an individual member than by characterizing cellular materials due to the latter’s sensitivity to a range of experimental artifacts such as size effects and strain rate sensitivity – both of which can be better controlled and studied at the level of an individual member [12].The member modeling approach, while promising, has its own set of challenges: it requires experimental characterization at the level of micro-struts or thin walls, which can prove to be difficult to accomplish due to the very fine feature sizes involved. However, this challenge is not insurmountable with proper experimental procedures and the right specimen design.

However, this challenge is not insurmountable with proper experimental procedures and the right specimen design. The second challenge is regarding the accuracy of the representation – how well can individual struts and walls describe what is happening in a complex cellular structure with different orientations and local thermal conditions that may be different from the struts is something that needs significant validation that may be specific to each process and material.CharacterizationEach of the three modeling methods mentioned previously relies on material properties from two sources – either this emerges from characterization of the bulk property by ASTM standard test procedures, or through testing of miniature specimens created to represent the member. In this section, we discuss the origins of these properties for both sources, prior to implementing them in the models for predicting behavior.

In this section, we discuss the origins of these properties for both sources, prior to implementing them in the models for predicting behavior. We begin though by establishing our experimental dataset that will form the reference for comparing each modeling method against. This data comes from prior work [1], the key aspects of which are recounted here.7961. Honeycomb CompressionIn prior work [1], an approach for extracting the elastic modulus of a honeycomb independent of its shape and size was proposed. As part of that work, square, hexagonal and triangular honeycombs of different sizes were tested under compression on an INSTRON 8801 tester. These honeycombs were manufactured with the Fused Deposition Modeling process on a Stratasys Fortus 400mc machine with ABS material. Details of the manufacturing and testing process, and the honeycomb design can be found in the cited reference. The primary result of interest is the compression load-displacement response shown in Figure 6.

The primary result of interest is the compression load-displacement response shown in Figure 6. The deformation followed a pattern shown in Figure 7. While Figure 6 is one representative sample, it adequately captures the elastic-plastic behavior we are trying to model and predict and as such is used as the comparative baseline in this study.Figure 6. Load-displacement response under compression associated with the honeycomb specimen shown in Figure 7Figure 7. Compression of a hexagonal honeycomb – images not to relative scale (details in [1])2. Bulk PropertiesBulk properties were obtained from the supplier datasheet [14] for the ABS-M30 material and the Fused Deposition Modeling (FDM) process that was used to manufacture the honeycombs used in this study. Since the datasheet specifies properties based on ASTM D638 standard specimen designs built in two different orientations, the orientation that best aligned with the honeycomb construction was selected (XZ axis).

Since we are interested in the elastic-plastic response, the tensile modulus and the yield strength were obtained from the datasheet, as 320,000 psi and 4,550 psi, respectively. Poisson’s ratio is not specified in the datasheet, so a value of 0.35 was assumed, following others in the literature [15].7973. Member CharacterizationTo replicate the cell walls in the hexagonal honeycomb, special specimens were designed that approximate the ASTM D638 specimens in the grip section, but the gauge section is replaced with a slender member, shown in Figure 8, that has the exact thickness (0.032 inch) and process parameters (exactly two contours – see reference [1]) as those used in each of the cell walls in the hexagonal honeycomb.Figure 8.

Specimen designed to have a thickness identical to the cell walls of the hexagonal honeycombSixteen of these specimens were manufactured on the same machine (Fortus 400mc) used to make the honeycomb, and carefully removed from the build platform without using any solvent, to be consistent with the processing conditions used in manufacturing the honeycomb. Six of these specimens were damaged during the removal process on account of their very small thickness, requiring delicate handling. The ten surviving specimens were tested under tension on the Instron 8801 with a 50kN load cell, relying on crosshead displacement for the computation of strains, and load cell readings for stress. The specimens were pulled at a displacement rate of 5 mm/min.The resulting stress-strain plots are shown in Figure 9. Failure strains vary for all samples, potentially on account of defects either during the printing process itself or specimen removal from the build sheet.

Failure strains vary for all samples, potentially on account of defects either during the printing process itself or specimen removal from the build sheet. With regard to the concern of this work, average elastic moduli was estimated at 285,972 psi. Conversion of the engineering stress-strain plots to true stress-strain plots, shown for a representative curve in Figure 10, generated a yield strength value of 4,855.3 psi and a tangent modulus of 6,317.1 psi associated with observed strain hardening. The yield point was estimated as the first point on the true stress-strain curve with zero-slope, in keeping with recommendations in the ASTM D638 standard [16]. Interestingly, while the modulus is lower than that quoted in the datasheet, the estimated yield strength is higher (the datasheet does not specify a tangent modulus).

Interestingly, while the modulus is lower than that quoted in the datasheet, the estimated yield strength is higher (the datasheet does not specify a tangent modulus). This may be attributable to dimensional variation as shown in Figure 3 – these were not explicitly corrected for, the assumption being made that the test specimen represents the honeycomb walls with regard to geometry and processing history. Table 1 summarizes these material properties from the two different sources.Figure 9. Engineering stress-strain plots from testing ten specimens under tension798Figure 10. Representative curve from Figure 9, replotted as a True Stress-True Strain curveTable 1.

Engineering stress-strain plots from testing ten specimens under tension798Figure 10. Representative curve from Figure 9, replotted as a True Stress-True Strain curveTable 1. FDM ABS Material Properties Used in ModelsBulk Property (Datasheet) Member CharacterizationYoung’s Modulus Es (psi) 320,000285,972Poisson’s Ratio νs0.350.35Yield Strength σys (psi)4,5504,855Tangent Modulus Ts (psi) Not Available6317.1Modeling MethodsOur goal with modeling is to predict the elastic and the initial plastic response of the honeycomb shown in Figure 7, based on the characterization data discussed in the previous section. There are two approaches we can take to develop these predictions: the analytical techniques developed by Gibson and Ashby [17] coupled with the use of homogenization techniques, treating the honeycomb as a block of solid material, or by using Finite Element Analysis (FEA).

For each of these methods, we can use material property data from different sources. We discuss each of these methods below.1. Homogenization ApproachThe analytical methods developed by Gibson and Ashby [17] can be used for modeling both the elastic modulus and the yield stress of a regular honeycomb structure. Specifically, the in-plane effective elastic modulus for a hexagonal honeycomb under compression is given as:(2)3𝑡𝑡 4 √3 𝐸𝐸𝑠𝑠 � 𝑙𝑙�1 2 1+(5.4+1.5𝜗𝜗𝑠𝑠)(𝑡𝑡/𝑙𝑙)∗ ∗ = 𝐸𝐸2 𝐸𝐸1 = * represent the elastic modulus in the two in-plane directions (which are equal * and E2 where E1 since the regular hexagonal honeycomb is a transversely isotropic structure), Es and νs are the799modulus and the Poisson’s ratio of the material that makes up the honeycomb. t and l are geometry parameters corresponding to the thickness and length of the hexagonal honeycomb unit cell.

t and l are geometry parameters corresponding to the thickness and length of the hexagonal honeycomb unit cell. Using this equation however tends to under-predict stiffness since it assumes the entire length l of the beam in the unit cell is contributing towards bending when in reality it is a smaller length that truly does so, as was described in a paper by Malek and Gibson that made a correction to the length calculation [18]. The revised calculation gives the reduced length lb as:(3)𝑡𝑡 2 cos 𝜃𝜃𝑙𝑙𝑏𝑏 = 𝑙𝑙 −where ϴ is 30o for a regular hexagonal honeycomb. This result for lb from equation 3 replaces l in equation 2.Gibson and Ashby also developed expressions for failure stresses in honeycomb depending on the mode of failure (brittle vs plastic collapse) [17].

Careful observation of the failure mode in the honeycomb experiments conducted in this study show the clear formation of a plastic hinge as shown in Figure 11, followed by interfacial separation, the latter of which is driven by the dual contour layout of the process. Our interest for this work is in the initiation of plastic collapse, for * which we may use the following equation derived by Gibson and Ashby for the plateau stress σpl honeycombs that exhibit such a failure mode:(4)22 3 �𝑡𝑡 𝑙𝑙𝑏𝑏�∗ 𝜎𝜎𝑝𝑝𝑙𝑙𝜎𝜎𝑦𝑦𝑠𝑠 where σys is the yield strength of the material.=To convert the stress and modulus equations into force-displacement diagrams that will enable a comparison to experimental data, we treat the honeycomb as a block of material with cross-sectional area and gauge length governed by the bounding box that the structure is contained by.

The cross-section area is estimated by the depth of the honeycomb (b) multiplied by the width (w), the gauge length for purposes of strain calculation was taken as the height of the honeycomb (h). These values for the honeycomb used in this study are given in Table 2.Figure 11. Close-up examination of deformation and failure mechanism of the honeycombs used in this study, clearly showing formation of a plastic hinge prior to interfacial failure800Table 2.

Close-up examination of deformation and failure mechanism of the honeycombs used in this study, clearly showing formation of a plastic hinge prior to interfacial failure800Table 2. Relevant dimensions of honeycomb used in this study [1]Thickness (t)0.032 inchLength (l)0.18 inchEffective Length (lb) 0.1615 incht/lb0.1981b 1 inchwh3.10 inches 3.15 inchesSubstituting the values in Tables 1 and 2 into equations 2 and 4 allows for the estimation of the following relationships between load and displacement in the elastic regime (elastic stiffness) and the load at yield, listed in Table 3 for each of the two material assumptions (from the datasheet and from the member characterization).Table 3.

Elastic and Yield Load-displacement Estimates Based on Homogenization ApproachBulk Property (Datasheet) Member CharacterizationElastic Stiffness (lbf/in) Load at Yield (lbf)4587.33 369.024099.52 393.76To relate this to the experimental result in Figure 6, the elastic stiffness value is used to provide an estimate of the initial force-displacement response up to the load at yield, at which point perfect plasticity is assumed (no hardening or softening) to create the graph in Figure 12. Both datasets do a reasonable job of predicting the elastic response at low strains but do not replicate the plastic behavior adequately. The bulk property estimation performs better in the elastic regime, but is worse in the plastic regime.Figure 12. Comparison between experimental data and homogenization model using bulk property and member characterization estimations2.

Comparison between experimental data and homogenization model using bulk property and member characterization estimations2. FEA ApproachUsing FEA we are able to explicitly model the honeycomb in its geometric detail, which we are unable to do with an analytical approach. The honeycomb CAD model was directly used801in ANSYS™ FEA software with one modification. Following our own observations and other published literature [19], it is known that fillets at the corners of honeycombs and lattices can significantly increase the stiffness of the cellular material. Data collected using blue-light scanning revealed an average fillet radius of 0.05 inches (shown for one specimen and corner in Figure 13). This corner was included in the CAD file for analysis to ensure greater accuracy in prediction of response of the honeycomb structure.Figure 13. (left) Blue light scan of a honeycomb structure enabling (right) the measurement of corner radius.

(left) Blue light scan of a honeycomb structure enabling (right) the measurement of corner radius. Results shown here are slightly higher than the average radius estimated for the hexagonal honeycomb at 0.05 inches (the speck on the scan is lost data associated with a marker)The model was setup as a 2D plane strain problem, with frictional contacts applied at the top and bottom platens (friction coefficient of 0.15). The precise value of this coefficient was not experimentally estimated – this number was selected since it was at the low end for what would yield a converging solution without rigid body motion. Displacements were applied at one platen, the other was fixed and used for the measurement of reaction forces. The model was setup for both material models: bulk property and member level characterization, from table 1. Figure 14 demonstrates the setup, along with a close-up showing the mesh.

The model was setup for both material models: bulk property and member level characterization, from table 1. Figure 14 demonstrates the setup, along with a close-up showing the mesh. Mesh refinement studies were conducted and once displacement results changed by less than 4%, further mesh refinement was ignored. Large deflection effects were included in the analysis.Figure 14. (left) 2D plane strain model with platens connected to honeycomb with frictional contacts and (right) close-up of an individual cell showing the mesh size as well as corner radius modeled after experimental measurements802Figure 15 shows results of the analysis, with equivalent plastic strains accumulating along two hinge points as observed experimentally and anticipated theoretically.

While the precise location of the hinging within the honeycomb is not the same, the formation of a collapsing series of cells along a 30 degree line is consistent – indeed this was observed experimentally for all specimens tested [1]. The reaction force – displacement plots for both material models is shown in Figure 16, comparing against the experimental dataset – the plastic region is captured with greater detail with the FEA model, as compared to the analytical homogenization approach, as is to be expected due to the local and nonlinear nature of the deformation.(a)(b)(c)(d)Figure 15. Equivalent plastic strain contour (a) showing localization of deformation consistent with experimental observations (b). Equivalent plastic strain contours (c) within a plastically deformed cell shows plastic hinging consistent with experimental observation (d) and with theory [17]803Figure 16.

Equivalent plastic strain contours (c) within a plastically deformed cell shows plastic hinging consistent with experimental observation (d) and with theory [17]803Figure 16. FEA model with two different material property assumptions compared against experimental resultLimitations of StudyThis study has several limitations that limit its immediate applicability. We point these outand suggest further work needed to expand this work.i. Sample size: Only one hexagonal honeycomb formed the basis of the experimental dataset. While it was part of the larger referenced study that looked at size effects [1], and fell within expected lines, it is by no means statistically significant. A larger sample size is needed to address this issue.ii. Geometry, Process and Material Specificity: This work was conducted with the Fused Deposition Modeling process using ABS and for the hexagonal honeycomb structure only.

Geometry, Process and Material Specificity: This work was conducted with the Fused Deposition Modeling process using ABS and for the hexagonal honeycomb structure only. For other combinations of process, material and geometry, findings may be different. Specifically, the formation of a plastic hinge is critical to the use of the analytical result as well as is likely a key reason why good agreement is obtained with an elastic-plastic material model (without accounting for fracture toughness, for example). Similar work needs to be conducted for different materials, geometries and processes to assess how these modeling methods hold up.iii. Strain Rate Dependence: This work ignores the effects of strain rate dependence, which is known to be significant for plastics at ambient temperatures and for metals at high homologous temperatures particularly.

Further, this may be exacerbated by the complex geometry and localization of strains in the structure, even if it is globally deformed at a seemingly quasi-static rate [20], [21]. More work is needed to assess the effect of strain rates on these modeling methods and the conclusions drawn below.iv. 2D vs 3D Modeling: This work was conducted for the relatively straightforward 2D geometry of the honeycomb. The estimation of member data for a 3D lattice can be significantly more challenging due to orientation effects. Additionally, computational expense is very significant for the analysis of such a 3D structure. In this case, a blended approach that combines member modeling with homogenization maybe more appropriate.804ConclusionsIn this work, we have compared two different approaches of modeling and two different characterization methods specifically for the Fused Deposition Modeling processing technology and for hexagonal honeycombs.

While all methods do reasonably well in the elastic regime, modeling plastic behavior requires incorporation of geometry effects in FEA. To enable a meaningful quantitative comparison between the curves all plotted together in Figure 17, we applied a correlation coefficient comparing each dataset against the experimental baseline. The correlation coefficient is given as:(5)∑(𝑥𝑥−𝑥𝑥̅)(𝑦𝑦−𝑦𝑦�) 22𝐶𝐶𝐶𝐶𝐶𝐶𝐶𝐶𝐶𝐶𝑙𝑙 (𝑋𝑋, 𝑌𝑌) =�∑(𝑥𝑥−𝑥𝑥̅)∑(𝑦𝑦−𝑦𝑦�)where x and y represent the two quantities being compared, and the barred quantities represent their respective averages. Using this correlation coefficient, we can now quantify how well representative each model is of the experimental dataset, and this is compiled for the four approaches in Table 4.

Using this correlation coefficient, we can now quantify how well representative each model is of the experimental dataset, and this is compiled for the four approaches in Table 4. This confirms our visual intuition, that the combination of a FEA modeling approach with material properties extracted from member level characterization most closely approximates the experimental data, with a correlation coefficient approaching 0.99.It is reiterated that the above conclusion applies only to the process, material and structure studied in this work. More work is needed to ascertain how this methodology would perform for other processes and geometries, and how it would work in 3D cellular structures such as lattices.Figure 17. A comparison of all four methods and the experimental dataset, replotted on one graphTable 4.

A comparison of all four methods and the experimental dataset, replotted on one graphTable 4. Comparison of Correlation Coefficients for the Four Models against Experimental Data Homogenization (Member) 0.9384MethodHomogenization (Bulk Property) 0.9595FEA (Bulk Property) 0.9210FEA (Member) 0.9857Correlation Coefficient805AcknowledgementsThis effort was performed through the National Center for Defense Manufacturing and Machining under the America Makes Program entitled “A Non-Empirical Predictive Model for Additively Manufactured Lattice Structures” and is based on research sponsored by the Air Force Research Laboratory under agreement number FA8650-12-2-7230. The U.S.

The U.S. Government is authorized to reproduce and distribute reprints for Governmental purposes notwithstanding any copyright notation thereon.DisclaimerThe views and conclusions contained in this document are those of the authors and should not be interpreted as necessarily representing the official policies, either expressed or implied, of the Government. Distribution authorized to U.S. Government Agencies and America Makes Members; Critical Technology. Other requests for this document shall be referred to AFRL/RXMS, Wright-Patterson Air Force Base, OH 45433-7750.References[1][2]T. Le, D. Bhate, J. M. Parsey, and K. H. Hsu, “Determination of a Shape and Size Independent Material Modulus for Honeycomb Structures in Additive Manufacturing,” Solid Free. Fabr. Symp., 2017. L. Liu, P. Kamm, F. García-Moreno, J. Banhart, and D.

viewed PaperCOMPARISON OF MULTIPLE HEAT TREATMENTS BY OBSERVING MECHANICAL PROPERTIES AND MICROSTRUCTURE OF LPBF FABRICATED ALUMINUM F357J. Diaz1,2, K. Caballero1,2, V.A. Medrano1,2, E. Arrieta1,2, M. Benedict3, D. Godfrey4, R.B. Wicker1,2, and F. Medina1,21 W.M. Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, TX 79968, USA  2 Department of Aerospace and Mechanical Engineering, The University of Texas at El Paso, TX 79968, USA 3 Air Force Research Laboratory, Wright-Patterson Air Force Base, Dayton, OH, 45433, USA 4 SLM Solutions Group AG, GermanyAbstractIn this paper, Aluminum F357 (AlSi7Mg), a material which is widely used in the automotive, aerospace, and additive manufacturing industries, will be analyzed after performing several heat treatments to enhance the properties of the material.

However, there is currently no standard for the usage and heat treating of F357 alloy; for that reason, ASTM F3318 standard will be followed for heat treating it. Having a comprehensive study on the performance of 3D-printed F357 benefits the automotive, military and aerospace industries due to the numerous casted components already in service and many becoming legacy components. This work presents mechanical and microstructural properties of F357 specimens fabricated with SLM technology and subjected to heat treatments; as-built, stress-relief, T6, hot isostatic pressing (HIP), and HIP+T6 heat treatments were applied. Furthermore, with the interest of the alloy performance in- service conditions, the specimens were subjected to artificial thermal aging.IntroductionAs additive manufacturing evolves into a more viable option for manufacturing components, more materials are being studied.

F357 is one of those materials that still needs to be studied more into depth so that it can be used where suitable since it has been mainly used for casting applications[1]. Aluminum F357 is anti-corrosive, and lightweight[2][3]. These characteristics make this material ideal for aerospace, military, and automobile applications[4].This study was performed using a Selective Laser Melting (SLM) M280 machine. SLM is a Laser Powder Bed Fusion (LPBF) AM technology which uses a laser beam to melt layers of metal powder to then form a part. This technology can build complex geometry components with good mechanical properties, comparable to the ones achieved with traditional manufacturing processes[5]. Aside from this, LPBF is also capable of delivering high density parts with high dimensional accuracy[6]. It is common to apply heat treatment to the manufactured samples in order to enhance the properties of the material[7].

It is common to apply heat treatment to the manufactured samples in order to enhance the properties of the material[7]. In general, heat treatments are a common post processing step for most metal AM technologies given that by applying a heat treatment, most of the residual stresses formed in the printing process can be released/eliminated as well as the porosity induced by either the printing process or the nature of the atomized powder[8]. In314particular, SLM can induce certain stresses due to the high temperature gradient while in the printing process[9]. Given that this material has no specific standard for heat treatment, ASTM F3318 – 18 standard was followed and several heat treatments were applied; as-built, stress-relief, T6, hot isostatic pressing (HIP), and HIP+T6. After applying these heat treatments, the specimens were aged at different temperatures and aging times (0hr, 100hr, 1000hr at 285 °F and 385 °F).

After applying these heat treatments, the specimens were aged at different temperatures and aging times (0hr, 100hr, 1000hr at 285 °F and 385 °F). Later they were tested and analyzed to observe their mechanical performance and microstructure.Materials and MethodsPowder FeedstockThe material used for this study is atomized F357 (AlSi7Mg) powder, obtained from Valimet, AM 357C. A Retsch Camsizer X2 machine was used to characterize this powder. This characterization process showed that the powder had a particle size distribution of D10: 24.4 μm, D50: 39.3 μm, and D90: 60.4 μm. Figures 1 & 2 show SEM images of the powder particles; as it can be seen, the particles seem to be mostly spherical.Figure1: SEM Powder Particles Figure 2: SEM Particles (Higher Magnification)Laser Powder Bed Fusion (LPBF)This technology additively manufactures components by melting layer by layer of metal powder with a laser.

This process occurs in an inert atmosphere such as in an argon gas environment which allows this to be a clean process. It is mostly suitable for rapid prototyping but can also be used for low & high volume production. As AM has grown during the past decades, this technology has evolved and become one of the primary technologies for metal AM and can now print a wide variety of materials. As the printing layer thickness is considerably small, the resulting surface finish of the printed components is of high quality. As mentioned above, LPBF has high dimensional accuracy and can produce highly dense parts. Although it is common for the parts to go through a heat treatment after printing due to residual stresses created in the process, this technology has shown it can deliver components with desirable microstructure and mechanical performance.

Figure 3 shows the machine in which all samples were printed.315Figure 3: SLM M280Heat Treatments and Thermal AgingAs mentioned before, several heat treatments were applied to the printed specimens (all asper ASTM F3318 – 18):SR1: parts shall be stress relieved or partially annealed in accordance with AMS 2771 except that the temperature shall be 285 °C (±14 °C), held for 120 min (15±) and cooled at a rate equal to air cooling or faster. Refer to AMS 2771 for soaking time and furnace class requirements.T6: parts shall be solution heat treated in accordance with AMS 2771 except the temperature shall be 530 °C (±6 °C) held for 360 min minimum, quenched in water or glycol as agreed upon by the part supplier and purchaser and aged at 160 °C (±6 °C) for 360 min minimum.

Refer to AMS 2771 for soaking time and furnace type requirements.HIP+T6: product shall be hot isostatically pressed (HIP) under an inert atmosphere at 100 MPa minimum within the range 510 °C - 520 °C, held at the selected temperature within ±14 °C for 180 ±60 min and cooled under inert atmosphere to below (93 °C). HIP should be followed be Condition T6 processing. Refer to AMS 2771 for soaking time and furnace class requirements.As-Built (NHT): no heat treat required. Parts are supplied in the as-built condition.In addition to the heat treatments, artificial thermal aging will be applied to the specimens to further analyze the behavior of the material and simulate its performance under service conditions.

The aging is divided into two temperatures, 285 °F and 350 °F, held at three different times (0 h, 100 h and 1000 h).Tensile Testing (Uniaxial Monotonic Tensile Load)After applying the proper heat treatment to samples, each of them were then machined into tensile coupons to then perform a tensile test as per ASTM E8 Standard Test Methods for Tension Testing of Metallic Materials. This test was performed in an MTS Landmark machine. When the316test was performed, this machine would obtain the mechanical properties such as Yield Strength, Elongation at Fracture, and Ultimate Tensile Strength among others.Figure 4: MTS Landmark; Tensile TestDensity MeasurementsA selected section was cut from the tested tensile coupons to then analyze it in an Accupyc II 1340 Pycnometer, shown in Figure 5, to obtain density measurements. The process involves gas displacement.

The process involves gas displacement. After weighing the sectioned piece, the density measurements can then be obtained by performing a calculation with the mass and volume values obtained from the weighing apparatus and pycnometer system respectively.Figure 5: Accupyc II 1340 PycnometerMicrostructure CharacterizationThe previous section used for density measurements was again cut in the X, Y, and Z plane direction to later be mounted in epoxy to then be grinded and polished with the use of an ATM SAPHIR 530 machine. Several pictures of the different planes were taken at multiple magnifications using an Olympus GX53 microscope.

Several pictures of the different planes were taken at multiple magnifications using an Olympus GX53 microscope. Later, the same mounted samples would be etched and analyzed using the same microscope to observe the melt pools.317Figure 6: Mounted Samples Figure 7: Grinding & Polishing SAPHIR 530Results and DiscussionMicrostructure AnalysisDensity for the material used in this study is 2.67 g/cm3; all density measurements takenshall be compared to this density.Process As-Built SR1 HIP T6 HIP+T6Density (%) 99.2 99.6 99.5 98.3 99.8Table 1: DensityAfter taking images of the as polished microstructure of the different samples, it can be observed that the as-built samples show some porosity and almost no sign of precipitates. Light precipitate density can be shown in the SR1 samples, with a lower level of porosity.

Light precipitate density can be shown in the SR1 samples, with a lower level of porosity. After observing the HIP, T6, and HIP+T6 samples, it can be seen that they are very similar; they show high levels of precipitates and small amount of porosity.Figure 8: As-Built X-Plane Figure9: SR1 X-Plane318Figure 10: HIP X-Plane Figure 11: T6 X-PlaneFigure 12: HIP + T6 X-PlaneMechanical PropertiesAfter the tensile tests performed, all data was analyzed to obtain the average Yield Stress, Ultimate Tensile Strength (UTS), Young’s Modulus, and Elongation at Fracture for all samples. As it can be seen in the table below, the samples that showed the best results on Yield Stress and UTS and Young’s Modulus were the as-built; horizontal and vertical orientations showed similar results. Nevertheless, as-built samples showed small percentage on Elongation at Fracture compared to the others, showing that they are brittle.

Most samples show enhanced performance on UTS and Yield Stress at 285 °F, specially at 100hr, and increased elongation as the temperature and aging time increase.320i s ki s k70 60 50 40 30 20 10 0605040302010000100hr 285 °FSLM AsBuiltSLM SR1SLM T6SLM HIPSLM HIP+T60.0020.0040.0060.0080.01in/inFigure 14: Stress vs Strain Graph 285 °F 100hr1000hr 285 °FSLM as-builtSLM SR1SLM T6SLM HIPSLM HIP+T60.0020.0040.006 in/in0.008Figure 15: Stress vs Strain Graph 285 °°F 1000hr3210.0120.01100hr 350 °FSLM as-builtSLM SR1SLM T6SLM HIPSLM HIP+T6504030i s k2010000.0020.0040.006 in/in0.0080.010.012Figure 16: Stress vs Strain Graph 385 °F 100hr1000hr 350 °FSLM as-builtSLM SR1SLM T6SLM HIPSLM HIP+T6504030i s k2010000.0050.01in/inFigure 17: Stress vs Strain Graph 385 °F 1000hrConclusionAll samples were built using as SLM machine, LPBF technology, and multiple heat treatments were applied to them.

This was done to further the research done on Aluminum F357 as it has no proper standards to heat treat this material. After applying the heat treatments, multiple322samples were subjected to thermal aging. Next, they were mechanically tested, and a microstructure analysis was performed. As it can be seen in the results section, this material showed high values in UTS and Yield Strength in its As-Built form, but showed higher levels of ductility as it was heat treated. This material also changes dramatically as it becomes aged, showing higher levels of ductility as it is aged with more temperature during longer times. Depending on the application, this study has demonstrated how the material’s properties can be changed/manipulated by heat treating it to obtain whatever characteristics are needed.Future WorkThis study is an ongoing research; several sections of this project are still being worked on such as hardness testing and etching the samples.

With these sections of the project being completed, the specimens will be further analyzed. After finishing these portions of the study, a better understanding of the microstructure and performance of this material under the different heat treatment conditions will be obtained.AcknowledgementsThe research described here was performed at The University of Texas at El Paso (UTEP) within the W.M. Keck Center for 3D Innovation (Keck Center). The views and opinions expressed in this article are those of the authors and do not necessarily reflect the official opinion or policy of any other agency.323Bibliography[1] K. Tsaknopoulos, J. Grubbs, M. Siopis, A. Nardi, and D. Cote, “Microstructure and Mechanical Property Evaluation of Aluminum F357 Powder for Cold Spray Applications,” Journal of Thermal Spray Technology, vol. 30, no. 3, pp. 643–654, Feb. 2021, doi: 10.1007/s11666-020-01140-4.[2] M.

COMPARISON OF POLYAMIDE 11 MECHANICALPROPERTIES BETWEEN LASER SINTERING AND TRADITIONAL MOLDINGDavid K. Leigh Harvest Technologies, Belton, TXABSTRACTTensile properties of laser sintering grade Polyamide-11 are processed using laser sintering, compression molding, and injection molding and the resultant mechanical properties are reported. The primary contributor to the enhanced mechanical properties of injection molded specimens is a fully healed polymer melt with preferred polymer chain orientation. It can be shown that laser sintering and compression molding specimens have comparable to compression molding specimens Ultimate Tensile Strength (UTS) and Elongation at Break (EOB).INTRODUCTION AND THEORY1.1: LASER SINTERINGAdditive manufacturing is a family of relatively new manufacturing processes utilizing integration techniques of digital solid models and the layer-by-layer addition of materials to create a three-dimensional solid.

Selective Laser Sintering®, developed at The University of Texas at Austin, utilizes three major steps to create the three-dimensional solid. 1. The first step of the process is the computational evaluation of a solid model and a subsequent creation of a set of cross-sectional vector fields referred to as “slices.” Each slice has a field of x-y vectors that represent the cross-sectional area of the model, with the aggregate total of these slices making up the complete 3-dimensional object. Each slice is prepared based upon a predetermined “layer thickness” which is typically 0.004 inches to 0.010 inches. Figure 1 illustrates the slicing algorithm process.Figure 1: Illustration of CAD model slicing and an individual 2-D Slice2. The second step of the process is to preheat a bed of powder, commonly a polyamide (nylon) material with an average particle size of 50 microns.

The second step of the process is to preheat a bed of powder, commonly a polyamide (nylon) material with an average particle size of 50 microns. Once the powder is heated to a sufficient temperature and the slice file is created, a directed energy beam (usually a CO2 laser) scans the vector field on the prepared bed of powder. The fused layer reproduces thecross-section generated by the vector field (or slice). A system schematic is illustrated in Figure 2.Figure 2: Schematic of Laser Sintering Process3. The final step is to add raw material to the process so that subsequent “slices” can be fused together. A layer of powder equal to the predetermined layer thickness that was used during the slicing process is added, heated, and subsequently fused to prior layers to create a solid physical model, illustrated in Figure 3.Figure 3: Photograph of powder bed being scanned.

Darker areas have been scanned by the laser while the lighter sections are unaffected powder.1.2: ADAPTATION OF LASER SINTERING FOR END-USE PARTSThe use of additive manufacturing and laser sintering in particular for low and medium volume production has become more important due to the cost savings associated with bypassing expensive tooling. (Ruffo, Tuck and Hague 2006) The laser sintering process has become more common in the production of end-use parts in various applications. This technique has allowed for streamlined design, decreased time to productize, minimal cost on engineering changes, and the ability to design for function rather than manufacturability. While there may be many benefits for certain applications, there are significant challenges that must be addressed prior to a broad acceptance of this technology.

While there may be many benefits for certain applications, there are significant challenges that must be addressed prior to a broad acceptance of this technology. The key challenge is the limited amount of theoretical orexperimental science as well as detailed materials databases specific for the use of laser sintering as a manufacturing process. The use of laser sintered parts for end-use production has increased the scrutiny of the process but has channeled corporate funding towards the end of creating a stronger knowledge-base.1.3: FORMATION OF LAYERS IN LASER SINTERINGFigure 4 is a graphical representation of the raw material powder bed. While the average particle size of the polyamide material is 50 microns, there is a distribution range of fine and coarse particles in a randomly packed bed. This powder bed is preheated to a point just below the melting temperature.

This powder bed is preheated to a point just below the melting temperature. If the particles are too hot they will pool, and there will be no distinction between the scanned area and the powder bed. If the particles are too cold, they will not have sufficient internal energy to allow for a full melt. In a region heated solely though laser energy will have significant dimensional distortion as the fused layer rapidly cools and shrinks due to the temperature gradient between the part and the surrounding powder bed. Part warping and in-build curl is created when the processed layer is fused to previous layers that have already undergone shrinkage.

Part warping and in-build curl is created when the processed layer is fused to previous layers that have already undergone shrinkage. This adhesion will create residual stresses as the processed layer shrinks at a different rate than the previous layers.Figure 4: Cross Sectional Illustration of a Laser Sintering Powder BedOnce the powder bed has been pre-heated, the laser will scan the subject cross-section using a field of x-y vectors at a prescribed speed, spacing, and laser power. These settings are optimized to yield an ideal melt pool such that there is sufficient energy to have a full melt without losing precision, seen in Figure 5. If the laser energy applied is too great, detail and precision are lost (analogous to writing on tissue paper with a large permanent marker).

If the laser energy applied is too great, detail and precision are lost (analogous to writing on tissue paper with a large permanent marker). There is a corresponding depth of penetration directly proportional to the total energy applied.Figure 5: Melt Pool Generated on the Laser Sintered Powder Bed1.4: PROCESS VARIABILITY AND PROCESS LIMITATIONSA significant limitation to any layer-based additive manufacturing process is the anisotropic nature of the process. In this case, a single layer (typically 0.004” thick) is fused to the previous layer. The properties within this layer are fairly consistent, but the bond between layers is typically viewed as weak. Intralayer bonding is achieved as the laser passes across the pre-heated powder causing particles to melt and flow together in the creation of a molten region defined by the slice.

Intralayer bonding is achieved as the laser passes across the pre-heated powder causing particles to melt and flow together in the creation of a molten region defined by the slice. Interlayer bonding is more challenging for three primary reasons: the previous layer that is being adhered to has had time to cool, the thin layer of powder that has been added serves as a thermal insulator and can inhibit laser penetration, and laser energy is absorbed and dissipated at a higher rate near the surface. This is discussed by Caulfield in “Dependence of mechanical properties of polyamide components on build parameters in the SLS process” and illustrated in Figure 6. (Caulfield, McHugh and Lohfeld 2006) This work showed fracture lines in the x-y plane in z-axis oriented specimens followed the layer lines where the bond was weaker between layers.Figure 6: Part-build orientations.

Black lines indicate layer interfaces.As a quality assurance step during the production of end-use parts, it is common to quantify build performance through the placement of standard ASTM D638 tensile specimens oriented in the x-axis (left to right), y-axis (front to back), and z-axis (vertical build direction). It is generally observed that there is little to no variation between the x and y axes and that the z- axis will yield a significantly lower elongation at break (EOB) than the x and y axis specimens.The primary cause of the lower mechanical properties in the z-axis can be attributed to layer-to-layer adhesion (interlayer bonding). Several papers have been dedicated to the characterization and optimization of the laser sintering process in relation to z-axis anisotropy.

Several papers have been dedicated to the characterization and optimization of the laser sintering process in relation to z-axis anisotropy. Several key factors that contribute to poor interlayer bonding (layer-to-layer adhesion): • Large layer size – The powder insulates the thermal energy and prevents the energy of thebeam from penetrating to the prior layer.Poor powder quality – Used powder will yield a high melt-flow and higher molecular weight. This is primarily caused through thermal aging and cross-linking of the polymer. The result of the high melt-flow is a polymer that is slow to melt with a high viscosity. The result is the creation of voids between layers.Poor thermal control – If the powder bed is processed in a way that the temperature is too low or is cooled at a high rate, it will inhibit the formation of a melt pool that penetrates the previous layer.

The result can be weak layer-to-layer adhesion as well as dimensional distortions.Low laser energy – High scan speed, low laser power, and large spacing between scan vectors (scan spacing) can lower the overall laser energy delivered to the powder bed. This decreased energy will not yield sufficient energy to penetrate the current layer and create strong layer-to-layer adhesion.1.5: PREVIOUS WORK ON PROCESS OPTIMIZATION AND CHARACTERIZATIONLaser sintering became a broadly used additive manufacturing process in the early 1990s with the introduction of Duraform® PA, a Polyamide 12 material. Christian Nelson published an article in 1995 speaking to the need for process and material optimization to improve part accuracy. (Nelson, McAlea and Gray 1995) The primary application of laser sintering in the early stages of development was functional prototypes and concept models with the chief competitor being 3D Systems™ Stereolithography (SL) process.

While laser sintering held an edge over the SL process in material properties, the SL had become the de facto standard in prototyping and held an edge in accuracy. The conclusions that Mr. Nelson made at the time were: • Use of better galvanometers to control the laser and scanners will increase accuracy andrepeatability.Optimization of scanning parameters can improve surface finish, accuracy, and repeatability. • Optimized powder particle distribution can yield surface finish comparable to those made with liquid based techniques.With increased utilization of the laser sintering process in the 1990s, there became a focus on process optimization and repeatability in regards to the use and re-use of raw material. Poor accuracy, surface finish, and overall part quality were seen in parts being made with Duraform® PA due in large part to inconsistent raw material and a thermal degradation of the material.

Poor accuracy, surface finish, and overall part quality were seen in parts being made with Duraform® PA due in large part to inconsistent raw material and a thermal degradation of the material. There were several schools of thought that were proposed to mitigate this issue:David Keith Leigh of Harvest Technologies developed work done to create a steady-state mixture of used and virgin based on quantified mechanical properties and qualified visual inspection of the parts. (Leigh and Gornet n.d.)Timothy J. Gornet of the University of Louisville developed work done on quantifying the melt flow rate and correlating this measurement to a usable standard. Virgin powder would be added to the powder feedstock based on a quantified melt flow reading. (Leigh and Gornet n.d.)John Choren proposed varying manufacturing parameters based on the powder age or thermal degradation. (Choren, et al.

(Leigh and Gornet n.d.)John Choren proposed varying manufacturing parameters based on the powder age or thermal degradation. (Choren, et al. 2001)Based on a paper written by Choren in 2001, it was noted that thermal processing parameters were stable but that laser power had to be increased over time. (Choren, et al. 2001) The mechanical properties achieved are shown in Figure 7 with the elongation at break absent from this data. In addition to this work, a 2008 paper by Yusoff showed the effects of powder degradation on surface finish. (Yusoff and Thomas 2008) These two published reports as well as work done by Leigh and Gornet illustrate the effect of powder melt flow (molten viscosity) on material properties, machine parameters, and part quality. (Leigh and Gornet n.d.)Figure 7: UTS Data taken from 2001 Choren powder life study. (Choren, et al.

(Leigh and Gornet n.d.)Figure 7: UTS Data taken from 2001 Choren powder life study. (Choren, et al. 2001)Further research in quantifying the effect of powder melt and mechanical properties was performed by Majewski at the University of Loughborough in 2008. (Majewski, Zarringhalam and Hopkinson 2008) This article was able to determine both visually and with Differential Scanning Calorimitry (DSC) that there were two distinct melt peaks representing areas in the processed part that had not been fully melted. It was proposed that the Degree of Particle Melt (DPM) had an overall impact on mechanical properties. A key conclusion in this research was stated:This work has shown that the percentage crystallinity of a two-phase SLS part, as calculated from a DSC chart, has an appreciable effect on the Tensile Strength and Elongation at Break, whereby a decrease in crystallinity leads to an increase in the mechanical property.

(Majewski, Zarringhalam and Hopkinson 2008)Figure 8: Laser Sintering Morphology taken from Particle Melt Study (Majewski, Zarringhalam and Hopkinson 2008)In work that has yet to be published at the University of Louisville, Thomas L. Starr and Timothy J. Gornet have research that further helps characterize the correlation of process settings, material morphology, and resultant mechanical properties. Starr correlates the amount of energy that is delivered to the part bed during laser sintering (Energy Density) to the resultant material properties. The formula for energy density is shown in Equation 1 and a plot of Yield Stress and Elongation at Break are in Figures 9 and 10, respectively. (Starr, et al. 2008) The energy to melt is shown to be insufficient for optimized material properties and it is proposed that there is an effective Energy to Melt Ratio. This excess energy is necessary for stronger layer-to-layer adhesion.

This excess energy is necessary for stronger layer-to-layer adhesion. In addition, Gornet shows in separate work that the use of microtoming is a practical tool in analyzing laser sintering morphology. (Gornet 2010) Figure 11 shows a microtomed cross-section of a specimen with layer lines being apparent in both magnifications.

Energy Density used from (Starr, et al. 2008).Figure 10: Plot of Elongation at Break vs. Energy Density used from (Starr, et al. 2008). There is no correlation between strength in Figure 9 that guarantees ductility.Figure 11: University of Louisville Laser Sintering Microtoming Sample (Gornet 2010)Based on the various research and characterization that has been done previously anddiscussed in this section, there are several key takeaways that will impact this research: • Intrinsic raw material properties have an effect on processing in laser sintering.

• The Degree of Particle melt study and the microtoming specimens appear to indicate amorphology that justifies weak layer-to-layer adhesion.Correlation of energy density to mechanical properties shows a strong correlation of energy density to tensile strength but with significant variability with respect to elongation at break.1.6: FRACTURE MECHANICSBased on the previous research done on the failure mechanisms in laser sintered polyamide (Starr, et al. 2008), the yield stress and ultimate tensile strength are relatively consistent among reported results and are comparable to reported material properties in other processes. The key deficiency in the laser sintering process is an elongation at break that is significantly lower than other manufacturing processes. If there is no significant degradation in the intrinsic material properties, this fast fracture of tensile specimens contributes to a decreased elongation.

If there is no significant degradation in the intrinsic material properties, this fast fracture of tensile specimens contributes to a decreased elongation. Fracture mechanics is the study of ductile materials failing due to brittle failure. There are three modes of failure with the primary failure mode in tension being Mode I, seen in Figure 12.Figure 12: Fracture Mechanics Failure Modes taken from (Twisp 2008).), of polyamide is significantly greater than what is observed in laser sintering while the fracture toughness remains the same. The correlation of fracture toughness to critical yield stress is shown in Equation 2. This correlation assumes that there is a crack within a part (a) or specimen and that the geometry of the crack and specimen can be quantified which produces a shape factor (Y).

This correlation assumes that there is a crack within a part (a) or specimen and that the geometry of the crack and specimen can be quantified which produces a shape factor (Y). A sampling of shape factors is shown in Figure 13.The critical yield stress, or fracture stress (σ#$%& (cid:12) ’()√+,(Equation 2)Figure 13: Fracture Mechanic Shape Factors (Roesler, Harders and Baeker n.d.)1.7: PROBLEM STATEMENTThis work will investigate the primary contributors to the mechanical failure of laser sintered nylon polyamide specimens. It is theorized that poor layer-to-layer adhesion, a primaryculprit in additive manufacturing anisotropy, can be attributed to intralayer porosity. The relative porosity of the interfacial boundary layers contributes to decreased ductility and fracture stress.EXPERIMENTAL PROCEDURE2.1: PRODUCTION OF LASER SINTERING SPECIMENSAll specimens in this study were built on machines owned and operated by Harvest Technologies in Belton, TX.

The specific machines used were 3D Systems™ sPro 60® or equivalent seen in Figure 14. These machines are the same as the base Sinterstation® 2500 with a 70 W CO2 laser, the addition of digital scanning, and active thermal control through the use of a black-body reference heat source. Samples were made in both PA 12 (Duraform® PA) and PA 11 (Duraform® EX and ALM FR-106).Figure 14: 3D Systems™ sPro 60® Laser Sintering MachineA total of 24 PA-12 specimens were produced and evaluated. These specimens were segregated based on the tensile results as described in Section 2.5, with selected samples being evaluated using SEM. After tensile testing, specimens were sputtered with gold particles for increased resolution and evaluated using SEM. Eight of the 24 specimens were used with redundant specimen data being discarded. Figure 15 shows an array of 8 tensile specimens with 4 being oriented primarily in the z-axis, and 4 being oriented primarily in the x-axis.

Figure 15 shows an array of 8 tensile specimens with 4 being oriented primarily in the z-axis, and 4 being oriented primarily in the x-axis. Three total builds were used with varying laser powers from 10% to 100% of optimized parameters to illustrate the full range of failure mechanisms typically seen.Figure 15: Tensile Specimen Location in Build2.2: MODELING INTERFACE BOUNDARY LAYER TENSION SPECIMENThe custom tension specimen used for characterizing the interface boundary layer was created in STL on an IBM PC with Magics 15 Rapid Prototyping Software from Materialise. The native ASTM D638 specimen was scaled and sectioned to insert a closed pack array of hemispheres to simulate a layer of powder, seen in Figure 16. The spheres were modeled at 0.100” diameter with the cross-section at the interface measuring 0.300”x0.500”.

The spheres were modeled at 0.100” diameter with the cross-section at the interface measuring 0.300”x0.500”. The half of the tensile bar that had a flat surface was combined with the specimen having the hemispheres to create a single tensile specimen STL file, seen in Figure 17. This operation was repeated for h/r Ratios of 0 to 1 at 10% increments, with h representing the depth of penetration of an intersecting plane with r, which represents the hemisphere’s radius. With the hemispheres measuring 0.050” radius, the interface was moved 0.005” for each subsequent sample. The h/r ratio dimensions are detailed in Table 3.

With the hemispheres measuring 0.050” radius, the interface was moved 0.005” for each subsequent sample. The h/r ratio dimensions are detailed in Table 3. Specimens were labeled with the software to qualify the resultant h/r ratio of each.Figure 16: Model of Close Packed Hemispheres on Tension Specimenr-hFigure 17: ASTM D638 Tension Specimen with Visible Interface Boundary2.3: PRODUCTION OF INTERFACE BOUNDARY LAYER SPECIMENSInterface boundary specimens were produced in optimized production builds and oriented primarily in the x-axis. Orientation of these specimens in the x-axis eliminates the possibility that layer-to-layer adhesion will impact test results. These tensile results were used to quantify the geometry effect of the interface boundary layer.

These tensile results were used to quantify the geometry effect of the interface boundary layer. Specimens were built in the same orientation as the x-axis specimens shown in Figure 15 using PA-11 (ALM FR-106) so that the testing could capitalize on the contrast afforded by a material that typically achieves greater than 50% elongation in the x-y plane.2.4: TENSION TESTINGTensile properties were measured using an MTS Insight 10 at Harvest Technologies, shown in Figure 19. The specimens produced were ASTM D638 specimens with a cross- sectional area at fracture of roughly 0.125” X 0.500”. The procedure for pulling the specimens was to use a 0.20 inch/minute rate with an extensometer that conforms to ASTM E83 class B2 with a range of at least 50%.

The procedure for pulling the specimens was to use a 0.20 inch/minute rate with an extensometer that conforms to ASTM E83 class B2 with a range of at least 50%. The extensometer is limited to 50%; thus, no quantifiable measurements exceeding that limit are recorded.2.5: SCANNING ELECTRON MICROSCOPY (SEM)The fracture surfaces were evaluated using a JEOL JSM 5610 Scanning Electron Microscope (SEM) at The University of Texas at Austin with a tungsten filament, 3 nm resolution, and magnification up to 300,000.Figure 18: MTS Insight 10, Extensometer, and Specimen2.6: CORRELATION OF PHYSICAL AND VISUAL PROPERTIESThe objective of this part of the research was to correlate the mechanical properties observed with visual inspection to obtain a better understanding of the failure mechanisms of laser sintered parts. The material was 3D Systems™ Duraform® PA, a Nylon 12 polyamide.

The material was 3D Systems™ Duraform® PA, a Nylon 12 polyamide. Tension specimens were added to a production build to evaluate the surface of a tension specimen and the fracture surfaces of both X and Z direction tension specimens. The process settings and build location of test specimens were such that three failure modes typical for laser sintering would be present for this experiment: 1. Delamination – A special type of brittle fracture that is characterized by very weak layer-to- layer adhesion. The fracture of a z-axis oriented specimen will break exclusively in the region between layers. Delaminated Z-direction specimens were processed at standard thermal parameters with laser power at less than 50% of recommended parameters.2. Brittle Fracture – Failure of a laser sintered specimen in tension within the elastic deformation range of a standard stress-strain curve. Fracture stress is less than yield stress with strains less than 10%.

Fracture stress is less than yield stress with strains less than 10%. Typical fracture surfaces have a 45o shear fracture.3. Ductile Fracture – This is the desired failure mode. A ductile fracture will yield nearly isotropic parameters. Processing parameters have been optimized for build quality with a sacrifice in efficiency and cost. There are two characterized ductile failure modes: a. Fracture Limited Ductility -- A fracture that occurs after the onset of strain hardening and typically before UTS and the onset of significant necking. Strains are typically 15-25%. b. Full Plasticity Failure – Fracture occurring after the onset of necking at UTS.

Strains are typically 15-25%. b. Full Plasticity Failure – Fracture occurring after the onset of necking at UTS. If this level of strain is reached, it is very common that failure will not occur until 40-60% strain.2.7: EVALUATION OF MELT FLOW RATE ON SPECIMEN FAILURE USING INJECTION MOLDINGIt has been observed that a primary contributor to poor mechanical properties of laser sintered components is the thermal aging, or degradation, of the material. Continuous exposure of the unsintered material to the elevated thermal environment causes cross-linking in the polymer chain and contributes to decreased physical and mechanical properties.

Continuous exposure of the unsintered material to the elevated thermal environment causes cross-linking in the polymer chain and contributes to decreased physical and mechanical properties. As the material undergoes thermal aging, the viscosity of the material is seen to increase and can be evaluated using a melt-flow measurement tool.A key consideration in evaluating specimen failure is whether this material degradation is an intrinsic material defect or whether this change in viscosity has an effect on process specific layer-to-layer adhesion. Unusable polyamide powder was segregated from production at Harvest Technologies and its viscosity (melt flow rate) evaluated.

Unusable polyamide powder was segregated from production at Harvest Technologies and its viscosity (melt flow rate) evaluated. This used polyamide powder was then processed in a BOY laboratory injection molding machine at The University of Texas to produce several tension coupons.2.8: MELT FLOW AND INJECTION MOLDING DEVICEPreviously used powder was measured in Dynisco Melt Flow Indexer at Harvest Technologies and then used as feedstock to run standard ASTM D638 tension specimens at The University of Texas. The melt flow indexer was used in conjunction with a drying oven, a calibrated balance scale, and Dynisco’s Polymer Test MIWORKS melt index test database in order to quantify the Melt Flow Index (MFI) of previously used powder compared to unused (virgin) powder. This device is designed to meet all of ASTM D1238 melt conditions. The BOY Laboratory injection molding machine was used to process the feedstock with standard PA settings using an ASTM D638 mold.

This device is designed to meet all of ASTM D1238 melt conditions. The BOY Laboratory injection molding machine was used to process the feedstock with standard PA settings using an ASTM D638 mold. Resultant tension bars were then tested at Harvest Technologies in the device described in section 2.4.2.9: COMPRESSION AND INJECTION MOLDING COMPARISONAn outside laboratory (Polyhedron Laboratories) was commission to make ASTM D638 tensile coupons in both compression molding and injection molding. A G-100T Model Morgan Press was used to create the injection molding specimens and processed at 450F with 40 hr conditioning. A PHI 50 ton press was used to create the compression molding specimens at 430F and 40 hr conditioning.

A PHI 50 ton press was used to create the compression molding specimens at 430F and 40 hr conditioning. The laboratory used an Instron Model 1114 tensile tester with an Electronic Instruments Research laser extensometer.2.10: GEOMETRIC CONSIDERATION OF THE INTERFACE BOUNDARY LAYERWhen laser sintered specimens are observed optically, there is a visible boundary layer between subsequently fused layers. This boundary layer is exclusively seen with X-Y plane orientation. This interface boundary layer appears to be a primary culprit in decreased mechanical properties and can be characterized as a layer of coplanar voids. Therefore, tensile specimens were created using CAD to simulate this boundary layer with varying degrees of severity. These specimens were then processed in the stronger X-direction to isolate the effect of the decreased z-direction mechanical properties on this experiment.Figure 17 illustrates the CAD model that was created to simulate coplanar powder particles.

The tension bar was scaled up from a standard ASTM D638 tension bar. The particle diameter was modeled at 0.100”.The two halves of the tensile bar were created independently to allow for a variation in the interface interference between a plane on one half and coplanar spheres on the other. This interface was defined by an h/r ratio seen in Figure 20. An h/r ratio (described in Section 2.2) of 0 had the spheres just touching the plane as illustrated in Figure 17. A ratio of 1 will have no effective interface boundary and is equivalent to a solid tensile bar. Eleven tensile specimens were produced with an h/r Ratio from 0 to 1 at an increment of 0.1. The material used was a polyamide 11 in an optimized production build. Optimized polyamide 11 production builds typically yield the highest mechanical properties and should provide the greatest contrast for the varying h/r ratio specimens.

Optimized polyamide 11 production builds typically yield the highest mechanical properties and should provide the greatest contrast for the varying h/r ratio specimens. This coplanar void creation is analogous to a “paper-towel” perforation with varying degrees of perforation or void density.Figure 19: Schematic of h/r ratioRESULTS3.1: TENSION SPECIMEN CHARACTERIZATIONThe tension testing performed yielded the results shown in Table 1. Data from 8 specimens are listed in this table and illustrated in the SEM photographs. Tension specimen nomenclature identifies the geometric location of the sample in the build as well as the build orientation. A “zx” specimen would be oriented primarily in the z-axis with the secondary axis oriented in the x-axis, consistent with ASTM F2921. An “x” or “y” specimen would be oriented flat in the x-y plane with the primary length along the x or y-axis, respectively.

An “x” or “y” specimen would be oriented flat in the x-y plane with the primary length along the x or y-axis, respectively. The x and y-axis specimens yielded better mechanical properties and more consistent results. The three primary failure mechanisms discussed in section 2.8 and enumerated in Table 1 can be seen in Fig. 21.Table 1: Tensile Properties and Fracture Mode of PA-12SampleZX 12.0.3 ZX 0.0.3 ZX 0.12.3 ZX 13.11.3 X 4.2.8 X 4.3.8 X 4.4.8Energy Density J/mm30.1 0.1 0.1 0.2 0.2 0.2 0.2Ultimate Tensile Strength (UTS) psi 494 2969 1278 5164 5886 5991 5763Elongation at Break (EOB) %1 1 1 3 7 6 7Fracture ModeDelamination Delamination Delamination Brittle Brittle Brittle Brittle+Y 1.4.144 1 Ductile fracture taken from independently optimized build.0.36099DuctileDelaminationBrittle FractureDuctile FractureFigure 20: Photographs of Tensile Specimen Failure Mechanisms. Specimens pictured are 1.50 inches in width.

Specimens pictured are 1.50 inches in width. The right specimen is 0.50 inches thick with the left two being 0.25 inches thick.3.2: MORPHOLOGY OF PA-12 TENSILE SPECIMEN FRACTURE SURFACESThe SEM image shown in Figure 22 shows an x-y sample with layering in the z-axis shown vertically. When inspected, the photographs show clear patterns for the layers with elongated horizontal voids between layers. Approximately energy density is approximately 0.10 J/mm3 with an energy-to-melt ratio of 1.0. Figure 23 is a magnified region of the same sample. It illustrates a single layer with voids both above and below the fully melted layer. There are also fully dense regions where there was sufficient energy to melt the layers together, creating a more homogeneous region.The surface of a Duraform PA (PA-12) XY tension specimen is illustrated in Figures 24 and 25. This shows the upper facing surface and the surface to which the loose powder must be bonded by the laser.

This shows the upper facing surface and the surface to which the loose powder must be bonded by the laser. The average particle size is roughly 50 microns and individual particles can be seen to be melted in Figure 25, appearing like a lily pad or pancake structure. Some particles can be seen to have melted with other particles while others appear to be isolated with voids between particles. Of particular interest in Figure 25 is the 250-m boulder looking particle in the center of the micrograph.The fracture surfaces of Duraform PA (PA-12) Z tensile specimens are shown in Figures 26 through 31. These specimens were evaluated with the camera parallel to the z-axis looking ata fracture generally in the x-y plane. Delamination is illustrated in Figures 26 and 27, brittle fracture in Figures 28 and 29, and ductile fracture in Figures 30 and 31. The surfaces photographed in Figures 26, 27, 30, and 31 were consistently parallel to the x-y plane.

The surfaces photographed in Figures 26, 27, 30, and 31 were consistently parallel to the x-y plane. The fracture surfaces photographed in Figures 28 and 29 were evaluating a 45° shear surface.3.2.1: Morphological ObservationsWhen evaluating the image of the fracture surface shown in Figure 22, a periodicity can be observed. With the understanding that the z-axis is the primary direction for added layers, a pattern of voids between layers is apparent. This stratification contributes to anisotropic material properties and significant weakness if a tensile stress is exerted parallel to the z-axis. In addition, these patterns of voids contribute to a significant reduction in elongation and ultimate tensile strength.When looking at the upper facing surface of the specimen shown in Figure 25, it appears that the 250-m particle was formed when several particles melted together but did not fully adhere to the surface.

When the particles cooled, there became a condensed grouping of several melted particles with a surrounding void. This voided region may contribute to voids in the part unless sufficient energy is used to re-melt these particles and bond with the surrounding surface. It is apparent from the analysis of these samples that the cracks propagate through voids created by a lack of full melt between layers. The general observations for each fracture mode are listed below:Delamination fracture surfaces can be seen in Figures 26 and 27. These images show a fracture of the Z specimen that is parallel to the x-y plane. Three distinct layers are represented as the crack propagation followed the area between two parallel layers and then migrated to neighboring layers. This pattern, when examined in Figure 27 shows that the particles are roughly 50 microns in diameter, which corresponds to the average particle size of the raw material.

This pattern, when examined in Figure 27 shows that the particles are roughly 50 microns in diameter, which corresponds to the average particle size of the raw material. When compared to the morphology of the top layer (Figures 24 and 25), it can be seen that the bottom layer has very little bonding to the previous layer.Brittle fracture surfaces can be seen in Figures 28 and 29. Voids are isolated and somewhat periodic, with similar size and spacing. Fracture surfaces show a ductile tear with transgranular fractures or river marks. These fracture surfaces are similar to the ductile specimen (Figure 30) but with void fraction on the failure surface higher.Ductile fracture surfaces can be seen in Figures 30 and 31. With an elongation of 44%, it may be observed that there is void nucleation where the material tore apart from itself. From Figure 31 shards of material and a very dense surface are observed.Figure 21: SEM of X Tensile Bar Fracture Surface.

(250x) PA-12, 0.2 J/mm3Figure 29: SEM of Y Tensile Bar - Ductile Fracture. (35x) PA-11, 0.3 J/mm3Figure 30: SEM of Y Tensile Bar - Ductile Fracture. (250x) PA-11, 0.3 J/mm33.3: INJECTION MOLDED TENSION SPECIMENSStandard injection molding parameters from a materials handbook were used to create fully dense ASTM D638 tensile specimens in-house in the Engineering Teaching Laboratory facility at The University of Texas at Austin. (ASM Engineering Materials Handbook n.d.) This particular machine was not in the greatest state of repair, but was sufficient for creating several tensile specimens from used PA-12 material. The powder was poured directly into the feed hopper of the injection molder and processed manually. The results of the tensile tests are presented in Table 2.

The powder was poured directly into the feed hopper of the injection molder and processed manually. The results of the tensile tests are presented in Table 2. Both specimens exceeded the range of the extensometer (50%) with one specimen pulled until it reached the mechanical limits of the load cell.Table 2: Injection Molded Tensile Results of Previously Used PA-12Sample 1 2(psi) 3884 3937./UTS (psi) 5572 psi 5530 psiEOB % >50% >200%Fracture Mode Ductile DuctileFigure 31: Photograph of Injection Molded Specimen3.4 EXTERNAL INJECTION AND COMPRESSION MOLDING SPECIMENSPolyhedron Laboratories, Inc. of Houston, TX provides plastics, polymers, and rubber testing. Laser sintering grade raw Nylon-11 material was provided to them for testing.

of Houston, TX provides plastics, polymers, and rubber testing. Laser sintering grade raw Nylon-11 material was provided to them for testing. The results reported can be seen in Table 3 below.3.5: H/R RATIO TENSION SPECIMENSTwo builds of 11 samples each were built and evaluated, with one set of samples having a finished surface and the other set having no post finishing done. There was no difference in ultimate tensile strength and elongation between builds. The data presented in Table 3 and Figure 32 are from the unfinished tensile specimens. There is a definite grouping of parameters for h/r ratios of 0.6 to 0.9 that corresponds with the strain hardening during necking of the specimens.It can be seen that all h/r ratio specimens followed the same stress-strain curve seen in Figure 33. The data collected was graphed in Figure 34 to illustrate the patterns observed in three primary regions of fracture.

Note that delamination is a special case of brittle fracture.Table 3: Injection and Compression Molded Tensile Results of PA-11Table 4: h/r Ratio Tensile Results of XY Oriented PA-11UTS (psi) 2954 3124 4021 4976 5312 5468 5882 6130 5994 6133 7391 * Specimen exceeded extensometer limit of 50%.Sample 1 2 3 4 5 6 7 8 9 10 11h/r Ratio 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0(psi) 2772 2854 3028 3056 3163 3047 2824 3090 3225 3214 3015./EOB % 4 4 6 7 9 9 17 17 16 18 >50*Figure 32: Stress-Strain curve for h/r Ration Specimens6050Full Plasticity Failurek a e r B t a n o i t a g n o l E %403020Brittle FractureFracture Limited Ductility10000.20.40.60.81h/r RatioFigure 33: Trends in h/r ratio tension properties3.5: X-RAY DIFFRACTIONAnother mode of decreased mechanical properties is seen in specimens that are near the build extent boundaries. Parts and specimens that are near the boundaries undergo a higher rate of cooling.

Parts and specimens that are near the boundaries undergo a higher rate of cooling. It has been observed in production builds at Harvest Technologies that the placement of specimens up to 0.5” deeper in a build yields better mechanical properties. Table 4 illustrates this with a clear increase in elongation with an increased depth that corresponds with a slower cooling rate.Table 5: Effect of Recrystalization on Tensile Properties of XY Oriented PA-11Specimen Depth (in)1 2 3 4.125 -.461 -.797 -1.133(psi) UTS (psi) EOB % 4757 2906 4749 2933 4876 2872 5007 2891./11 12 14 15Figure 34: Stress-Strain curve for x-ray diffraction specimensIt was theorized that the rate of cooling could correspond to a degree of crystallinity in the resultant specimens. Figure 36 shows x-ray diffraction results for four specimens listed in Table 5 and Figure 35.

Figure 36 shows x-ray diffraction results for four specimens listed in Table 5 and Figure 35. The periodicity in the x-ray diffraction figure is a function of the base material and the width of the peaks and valleys represent the level of crystallinity of the material. When observing these plots, periodicity and slopes appear equivalent. Based on these results, there appears to be no functional difference in crystallinity that could contribute to an increase or decrease in mechanical properties. Therefore, it is more likely that there is an optimal thermal processing dwell time above the recrystallization point promoting a stronger inter-layer adhesion.Figure 35: X-ray diffraction results showing equivalent periodicity and crystallinity.DISCUSSION4.1: MORPHOLOGY OF PA-12 TENSILE SPECIMEN FRACTURE SURFACESIt is apparent that the material properties observed in Table 1 can be correlated to the extent of voids present between sintered layers.

With very little adhesion between layers,delamination and elongation values near 1% are seen. With sufficient energy to melt through the target layer into the previous layer, a fully dense part will yield elongations well above 10%. As the layer has a higher degree of melt, the voids will decrease through the brittle range until the voids no longer contribute to the fracture.Figures 22 and 23 are photographs of specimen voids being created between subsequently built layers through the lack of complete particle melt to the previous layer. Figure 37 shows a larger separation between layers and represents the sample seen in Figures 26 and 27 (delamination). Since the voids are coplanar, they are analogous to pre-existing cracks. The crack propagation in this example will cause a shear between layers resulting in delamination.Figure 38 represents the sample seen in Figures 28 and 29 (brittle fracture) with voids present, but with the voids being isolated and periodic.

The crack propagation in this example is much less directional since the voids are no longer exclusively coplanar. A fracture may originate between specific layers, but will typically travel through voids in neighboring layers since the distribution of voids is much less ordered.Figure 36: Lightly adhered layers contributing to delamination.Figure 37: Incomplete fusion contributing to brittle fracture.Figure 38: Initial Layer of Sintered PowderFigure 39 is a graphical representation of an initial layer created in the laser sintering process. The depth of this layer is dependent upon the laser energy, but regardless of the amount of energy on the first layer, the downward facing surface will be defined by the raw material particle size as seen in Figure 26. The top of the layer is flat and will look like the surfaces seen in Figure 24 and 25.

The top of the layer is flat and will look like the surfaces seen in Figure 24 and 25. This irregular downward facing surface is the impetus for void creation.To insure the best properties of a part created though the laser sintering process, layers must be processed in such a way as to minimize or eliminate voids between layers. Voided regions between layers define the fracture mode. The h/r ratio experiment discussed in section 2.11 is illustrated again in Figure 40. The interface boundary layer created with the h/r ratioshown in Figure 39 illustrates the close proximity to the coplanar voids illustrated in Figures 37 and 38.Figure 39: Simulated interface boundary ratio using an h/r ratio4.2: EFFECT OF REUSED POWDER ON PARTICLE MELT IN INJECTION MOLDING AND COMPARISON OF LASER SINTERING TO COMPRESSION MOLDINGBased on the tensile data acquired, it was apparent that degraded powder did not yield poor mechanical properties when injection molded.

This test provides us a good degree of confidence that the degradation of mechanical properties in the laser sintering process is not an intrinsic material property defect but a property that affects the processing of the polyamide material in laser sintering. The melt flow index (MFI) of new powder is generally a value of 30g/10sec which corresponds to a low viscosity and relatively high melt flow rate. An MFI of 1g/10sec is a very low flow rate with a high viscosity. The material used in the injection molding tension specimens was previously used with an MFI of 1g/10sec.When comparing the mechanical properties measured in laser sintering, it is apparent that laser sintering is comparable to compression molding while exhibiting limited ductility when compared to injection molding. The fact that compression molding and laser sintering do not have the benefit of shear and complete mixing contributes to weaker specimens due to polymer melt surfaces that are not sufficiently healed.

There can also be a presence of voids in these melts that is not typically seen in injection molding which can contribute to fast fracture.4.3: COPLANAR VOID (H/R RATIO) BEHAVIORThe existence of an interface boundary layer can be illustrated in Figures 22 and 23 from the SEM images. In addition, work done at the University of Louisville with microtoming and optical microscopy of samples shows evidence of a boundary region between layers. The work cited here and in Figure 11 is not published at the time of this writing. (Gornet 2010)The images obtained from the University of Louisville show a cross section in a sample that is oriented the same as the fractures shown in Figure 22. The samples were created by using a microtoming technique common in medical pathology. The samples were sliced in layers roughly 5 microns thick to evaluate individual layers.

The samples were created by using a microtoming technique common in medical pathology. The samples were sliced in layers roughly 5 microns thick to evaluate individual layers. It can be seen in Figure 11 that the fused layers (lighter color) are fully dense and the particles at the bottom of the layer (darker) are nylon particles that were not sintered. It can also be observed that there are several tears in the sample shown in Figure 11 that correspond with these darker (unmelted) regions.When analyzing the graph in Figure 34, there are two distinct trends present with an outlier at an h/r ratio of 1.0. An h/r ratio of 0 to 0.5 represents fracture during elastic strain and before strain hardening. There is also a grouping of these regions observed in the stress-strain curve in Figure 33.The use of an h/r ratio helps quantify the failure modes in a macro scale and can help in identifying material process improvements that may help on a micro scale.

This correlation is significant.4.3: PART DENSITY OR CRITICAL CRACK SIZEOne metric that is commonly used to evaluate laser sintering parts is the density of the specimen. The boundary layer density was calculated using a slice taken from the CAD file of the periodic layer and interface boundary (r-h). The CAD file volumetric density data is plotted with the corresponding h/r ratio in Figure 41. This figure shows the density of a single layer based on the corresponding h/r ratio. If layers are consistent throughout the build, the resultant part density would be the same as an individual layer.

If layers are consistent throughout the build, the resultant part density would be the same as an individual layer. The issue with this assumption is that the mechanical properties do not show the same trend when comparing the elongation data in Figure 34.y t i s n e D r e y a L y r a d n u o B1.110.90.80.70.60.5y = -0.2318x2 + 0.4421x + 0.78600.51h/r RatioFigure 40: Correlation of effective part density and h/r ratio.It is possible that the jump in elongation data shown in Figure 34 when going from an h/r ratio of 0.9 to 1.0 is that the effective crack size hits a critical value. The critical crack size can be calculated using the formula shown in Equation 2 and reduced to Equation 3. To calculate the smallest critical crack size (ac), it is assumed that there is a simi-infinite plate and a single crack, using a value of 1 as a shape factor (Y). The critical crack size is assumed to be critical at the point the plate would yield vs.

The critical crack size is assumed to be critical at the point the plate would yield vs. fail due to fast fracture. Therefore, the yield stress would be used as the fracture stress. The data and results for the critical crack size calculations for injection molded polyamide are shown in Table 6.(cid:14)(cid:23) (cid:12)0 1 234(cid:23) 5679 8(Equation 3)Table 6: Theoretic Critical Crack Size of PA-11 and PA-12 (CES Edupack n.d.)Material PA-11 PA-12./(psi)8300-8700 3000-6110√=>) (psi :;< 3330-7310 1840-5530Y 1 1ac (in) - min 0.051 0.1197When evaluating the relative porosity of the interface boundary layer and disregarding the porosity as a whole, more intuitive results are obtained. The failure of any specimen in additive manufacturing will happen at that one layer or layers that are the most porous locally (largest ac).

The failure of any specimen in additive manufacturing will happen at that one layer or layers that are the most porous locally (largest ac). Given that fact, the relative porosity of the boundary layer in the h/r ratioexperiment was calculated and the results were plotted in Figure 42. The general form of the equation given by Haynes is shown in Equation 4 where ε is the relative porosity and C is a constant. (Haynes 1977)Relative Elongation = 	(cid:21)0?@(cid:28)⅓ B0?C@9D½(Equation 4)0.160.14n o i t a g n o l E e v i t a e Rl0.120.10.080.060.04C = 4,0000.02000.20.40.6Relative PorosityFigure 41: Relative porosity of interface boundary layer vs. relative elongation.When correlating fracture stress to the morphology observed, it is apparent that there is sufficient intralayer porosity for one or more critical cracks exist in the specimen prior to loading in the “delamination” examples shown in Figures 33 and 34.

As the specimen undergoes a tensile load and starts to elongate, fast fracture occurs due to a pre-existing crack equal to or greater than the critical crack size (ac), governed by Equation 3.The intrinsic cracks in samples that fail after the onset of strain hardening (again seen in Figures 33 and 34) are less than ac but are of sufficient density to inhibit the onset of significant necking. For those specimens that fail at strains seen after the onset of necking at UTS the relative porosity is greatly reduced. It is very rare to see z-axis strains approach 50% as shown in Figure 34 for the h/r ratio of 1 due to the intralayer porosity.CONCLUSIONSLaser sintered specimens are comparable to compression molding specimens in modulus, ultimate tensile strength, and elongation at break.

Laser sintering exhibits a significantly reduced elongation at break when compared to injection molding.Intralayer porosity is responsible for decreased mechanical properties in laser sintered polyamide. Ductility measured in the z-axis is particularly susceptible to the degree of porosity seen between layers and tends to segregate failure modes of specimens into three broad categories: brittle fracture, fracture limited ductility, and full plasticity failure.Thermally degraded powder does not affect the tensile properties when used in the injection molding process. These degraded powders poor (low) viscosity. Poor viscositypowders negatively affect the laser sintering process. It is assumed that the high viscosity of thermally degraded powders contribute to poor layer-to-layer adhesion.Poor layer-to-layer adhesion contributes to the formation of an interface boundary layer. This is evident in both SEM and optical microscopy.

This is evident in both SEM and optical microscopy. Based on the experiments run, it can be concluded that the primary cause of the lower mechanical properties in the z-axis can be attributed to layer-to-layer adhesion and the geometric characteristics of the interface boundary layer.There are three distinct modes of failure in laser sintered specimens: brittle fracture, fracture limited ductility, and full plasticity failure. It is apparent from the analysis of these samples that the cracks propagate through voids created by a lack of full melt between layers.The use of an h/r ratio and the relative density of the interface boundary layer help qualify the regions of fracture. An h/r ratio of 0.6 demonstrates a shift between brittle and ductile fracture modes. When the h/r ratio approaches 1.0, there is a major jump in mechanical properties when the interface boundary layer is significantly minimized.

When the h/r ratio approaches 1.0, there is a major jump in mechanical properties when the interface boundary layer is significantly minimized. This phenomenon may correlate to a unique coplanar shape factor or a critical crack size that instigates fast fracture.More work is necessary to understand the lack of layer-to-layer adhesion in laser sintered polyamides. Similar materials when molded will have elongation at break (EOB) values of 200- 400% while the laser sintered specimens rarely exceed 50%. There is a need to quantify the degree of crystallinity, the existence of micro-cracks or voids that correspond to a critical crack size, the fracture toughness of laser sintered polyamide, and the degree of cross-linking between powder particles in the melt pool.

These further studies may allow insight into unique polymer chemistry that will aid in the processing of laser sintering polymers and yield a higher degree of confidence in end-use laser sintered parts.ReferencesASM Engineering Materials Handbook. Vol. Plastics.Caulfield, B, P.E. McHugh, and S Lohfeld. "Dependence of mechanical properties of polyamide components on build parameters in the SLS process." Journal of Material Processing Technology (National Centre for Biomedical Engineering Science, Department of Mechanical and Biomedical Engineering, National University of Ireland) 182, no. 2007 (September 2006): 477-488.CES Edupack, Software. Edited by M. Ashby. Granta Material Intelligence . GrantaDesign Limited.Choren, J, V Gervasi, T Herman, S Kamara, and J Mitchell. "SLS Powder Life Study."SFF Symposium. Austin, TX: The University of Texas at Austin, 2001. 39-45.Gornet, T, interview by D.K. Leigh.

Conventional arc welding requires a maximum interpass temperature to limit any reduction in mechanical properties, but this may not be practicable for wWA-DED. In this study, two interpass process control methods, one with maximum interpass surface temperature and the other with constant dwell time, were adopted to deposit low alloyed steel walls while maintaining the same feedstock and heat input values. Thermocouples were inserted at three different positions in the walls during deposition, to record the thermal profiles. Test samples extracted from walls exhibited similar tensile strength (~10 MPa difference) and hardness values. Microstructural evaluation showed the presence of interlayer regions with alternating coarse and fine bands of ferrite grains, irrespective of the interpass control method.

Microstructural evaluation showed the presence of interlayer regions with alternating coarse and fine bands of ferrite grains, irrespective of the interpass control method. These findings suggest that dwell time control is better for productivity.Keywords: Wire-arc DED, Low alloyed steel, interpass surface temperature, interpass dwell time, thermo-couple measurement.IntroductionWire-Arc Direct Energy Deposition (WA-DED) is an emerging metal Additive Manufacturing (AM) method that possesses the potential to revolutionize the way large metal parts are produced. Like other AM methods, WA-DED introduces a pool of liquified metal from the wire feedstock by the electric arc that is deposited upon a substrate to fabricate a 3- dimensional component. In comparison with other metal AM processes, WA-DED has higher deposition rates (15 - 160g/min) [1], the ability to fabricate structural components with shorter lead time, and high material usage efficiency [1].

Compared to laser based and electron beam based AM methods, the arc imparts high heat input during deposition. Combining the high heat input with repetitive deposition subjects the component to multiple reheating cycles. This gives rise to complex thermal cycles across different parts of the build contributing towards anisotropy in mechanical properties and microstructure [2,3]. Hence, the WA-DED process efficiency and material performance (microstructure and properties) of the as-deposited material are directly influenced by processing parameters, control methods, and feedstock composition [1].

Thus, extensive research has been made to optimise the processing parameters such as current, voltage, torch travel speed, wire feed speed, contact tip to work distance (CTWD), torch angle [4-6], and attempts have been made to introduce process control methods such as maintaining a constant interpass surface temperature [7] and dwell time [8, 9].1489Given the considerable research work pursued in optimizing these conditions, less work has been done on establishing a relationship between process parameters, control method (interpass surface temperature and dwell time control), resulting thermal cycle and its impact on material microstructure and mechanical properties. This interdependent relationship directly affects the microstructural evolution and mechanical properties in the as-deposited condition. Measurement of thermal profile during deposition makes it possible to establish this correlation.

Measurement of thermal profile during deposition makes it possible to establish this correlation. In this work, in-situ thermal measurements with K-type thermocouples have been carried out during wide wall deposition to understand the effect of process parameters on peak temperature and cooling rates. The influence on interpass control methods on the thermal profile is also done for the purpose of comparison.Experimental methodsIn the current study, two thick wall depositions of low carbon steel were fabricated by WA-DED system (as shown in Figure 1(a)), with in-situ temperature measurements made using K-type thermo-couples embedded within the wall during deposition. Two interpass process control methods were used for deposition of the walls, i.e., maintaining maximum (i) surface temperature (250°C) and (ii) dwell time (200s). Thermo-couples were placed at three different regions of the wall during deposition, for in-situ temperature measurements.

Thermo-couples were placed at three different regions of the wall during deposition, for in-situ temperature measurements. The thermal history obtained includes local cooling rates experienced and peak temperature attained during subsequent reheating stages (i.e., next layers).The chemical composition of the wire feedstock obtained from ESAB, is given in Table 1. The WA-DED system comprises a Fronius CMT advanced 4000R welding power source with a FANUC ARC Mate 120iC six-axis industrial robotic arm for housing and automated welding torch movement. The shielding gas used during the deposition process is M26 gas [10] along with an extraction unit. The heat input was set at a constant value of 2.4 kJ/mm by selecting the process parameters such as current (I), Voltage (V), travel speed (v) and wire feed speed (WFS) as given in Table 2. Heat input values were constantly recorded through the process using a AMV 4000 welding monitor (Triton Electronics Ltd, UK).

Heat input values were constantly recorded through the process using a AMV 4000 welding monitor (Triton Electronics Ltd, UK). For temperature measurement, calibrated K-type thermocouples of wire diameter 1mm, were used. The thermo- couples were ceramic insulated to guard from overheating by the electric arc during subsequent deposition. The overall time taken for depositing two walls were approximately 9 hours (for W1) and 7 hours (for W2), respectively. Different samples extracted for mechanical testing are specified in Table 3.1490Figure 1: (a) Wire-Arc DED setup for low carbon steel wall deposition.

The substrate surface was deburred with a steel wire brush and cleaned with acetone before the deposition. To obtain uniform heat distribution profile for a layer, oscillated or weave strategy was adopted where the welding torch travels across the width of the layer in an oscillating manner (as shown in Figure 1b).1491A total of 59 layers were deposited to reach the wall height of 203 mm with thickness 26 mm and length 325 mm. The regions of thermocouple insertion (TC-M1 to TC-M3) are shown in Figure 2. After deposition, the walls were sectioned from the substrate and surface grinded to individual plates which were non-destructively tested with X-ray radiography to identify regions of macro-defects and thermo-couple insertion. Tensile, hardness and charpy impact test samples were extracted from the as-deposited walls as shown in Figure 2.

Tensile, hardness and charpy impact test samples were extracted from the as-deposited walls as shown in Figure 2. The specimens were extracted across two different orientations, vertical (parallel to the build direction) and horizontal direction (perpendicular to the build direction).Figure 2: Specimen extraction plan for mechanical testing (tensile, hardness and Charpy) and microstructural analysis. M1, M2 and M3 denote the thermo-couple positions.Results and DiscussionThe in-situ measurements made by the thermo-couples placed at three different regions of the wall, M1 (top), M2 (mid), and M3 (bottom) are given in the Figure 3. The maximum range of the thermo-couple measurement is 1400°C which makes it difficult to measure the thermal cycle for the first pass, as the temperature under the arc reaches as high as 2600°C.

The maximum range of the thermo-couple measurement is 1400°C which makes it difficult to measure the thermal cycle for the first pass, as the temperature under the arc reaches as high as 2600°C. Hence, temperature recording was resumed from the second pass of deposition, and the cooling rate values were calculated from passes that attain full austenisation during reheating.Table 4: Summary of time taken to cool from 800C to 500C (t8/5) and the cooling rates from regions M1-M3 for W1 and W2.W1W2SampleM1M2M3M1M2t8/5 (s)58.760.864.211090Cooling rate (°C/s)5.114.934.672.733.3The thermal history experienced at different regions of deposition along with their corresponding cooling rates and t8/5 is given in Figure 3 and Table 4 respectively. The thermal history from the bottom region of W2 is not included due to thermo-couple detachment from1492the wall, after second pass of deposition, leaving with only one recorded pass of deposition.

The cooling rates are calculated from the time taken for the layer to cool down from 800°C to 500°C, t8/5, similar to calculations made for weld temperature measurements. The t8/5 value for all measurements was calculated from the third deposition pass except for the middle region measurement of W2 (W2.M2) since it did not undergo full re-austenisation after the second deposition. In comparison, W2 had a slower cooling rate which is due to heat accumulation, also evident from the increasing interpass surface temperature (from Figure 3).Figure 3: Temperature measurements from different regions of the as-deposited wall. M1 - Top region, M2 - Middle region and M3 - Bottom regionMicrostructural analysisMetallographic samples were extracted 3 cm away from the thermo-couple inserted region (M1-M3) to avoid any compositional change due to diffusion from thermo-couples.

Since the samples are extracted closer to the thermocouples, they were considered to possess the sample thermal profile as M1-M3 regions and referred with the same nomenclature. The samples were mounted in Bakelite (as shown in Figure 4(a)), then grinded and polished with different silicon carbide polishing papers followed by polishing with 3µm, 1µm, and 0.25µm diamond suspensions. The preparation was finished by polishing with 40nm silica suspensions for optical microscopy and Electron Back Scattered Diffraction (EBSD) analysis. Initial analysis commenced with light optical microscopy, after which grain size determination was continued with EBSD analysis. Finally, the samples were tested for hardness values after microscopy analysis.1493Figure 4: (a) W1 -Macrophotograph. (b)shows the interpass regions with alternating bands of coarse and fine ferrite grains.

(b)shows the interpass regions with alternating bands of coarse and fine ferrite grains. Optical micrographs of (c) fine grained and (d) coarse grained regionUpon etching with 2% Nital, it was observed from macro-photographs (Figure 4(a)), the presence of interlayer regions corresponding to layer height for each pass of deposition, for both the walls. The interlayer regions were present for both the walls, W1 and W2, across the deposition height. For the sake of reporting, one metallographic sample is chosen (W1.M1) and its microscopy results are shown in this section, while the quantitative results were obtained by performing the same analysis on all these samples. For W1.M1, the mounted and etched samples were viewed under optical microscope to reveal the presence of coarse and fine grained regions in the interlayers (Figure 4(b)).

For W1.M1, the mounted and etched samples were viewed under optical microscope to reveal the presence of coarse and fine grained regions in the interlayers (Figure 4(b)). Upon viewing at higher magnification, the difference in grain size is evident, with regions of fine and coarse equiaxed ferrite grains, as shown in Figure 4(c) and Figure 4(d), respectively. For W1.M1, the Inverse Pole Figure (IPF) map is shown in Figure 5.To quantify this difference in grain size and its distribution, further microstructural analysis was conducted with an Oxford Instruments NordlysMax EBSD detector fitted on a Zeiss Sigma Scanning Electron Microscope (SEM). The step size chosen for EBSD analysis was 1 µm while the grains were analysed with minimum detection area of 10 pixels (10 µm2) and a critical misorientation of 5°. Since the width of these interlayer regions varied from 3 to 3.5mm, automated scans across the build direction were done to construct a montage area across the interlayer region.

Since the width of these interlayer regions varied from 3 to 3.5mm, automated scans across the build direction were done to construct a montage area across the interlayer region. Again, to demonstrate the difference in grain size and distribution between these regions, sample W1.M1 is chosen for representation. From these scans, coarse and fine-grained regions were highlighted as different subsets, as shown in Figure 5, to estimate the average grain size and its distribution. The IPF map exhibits lack of any preferred crystallographic orientation in the microstructure, which is commonly reported from other AM processes [11] and wire-DED deposition of other alloys [12].

The IPF map exhibits lack of any preferred crystallographic orientation in the microstructure, which is commonly reported from other AM processes [11] and wire-DED deposition of other alloys [12]. Lack of preferred orientation was observed from EBSD analysis of other samples too.1494Figure 5: (a) EBSD Inverse Pole Figure (IPF) maps scanned across the interpass region of W1.M1 position to analyse the (b) coarse and (c) fine grain regions along with grain size distribution.Determination of grain size and distribution were made according to the previously mentioned criteria. The average grain size measured from the coarse and fine-grained regions were 6.43µm and 4.62µm, respectively.

The average grain size measured from the coarse and fine-grained regions were 6.43µm and 4.62µm, respectively. Given the difference in average grain size is less significant, further EBSD analyses were analogously made across the interpass region and grain statistics were determined from the whole scan, instead of dividing the scanned region into different subsets (coarse and fine grained regions).Rest of the microscopy samples were analysed by wide montaged EBSD scans to obtain more grain information (like W1.M1 analysis). The average grain size values and their distribution, for different microscopy samples, are given in Table 5 and Figure 6, respectively. Despite the difference in thermal profile and cooling rate (~3°C/s between W2 and W1) the grain size distribution remains similar across the deposition for W1, while for W2, as the deposition progresses slight coarsening is evident.

This can be observed by the variation in grain size across build direction (i.e., M1 to M3) for W2.Table 5: Average grain size determined from montaged EBSD scans for different metallographic samples.SampleW1W2ni a r g e g a r e v A)m µ (e z i sM1M2M310.04 ± 5.310.49 ± 5.810.5 ± 5.711.26 ± 5.612.24 ± 5.810.5 ± 5.51495Figure 6: Average grain size distribution analysed from EBSD scans across different location (M1 - M3)Following EBSD analysis, the prepared samples (extracted along the wall height, Figure 2) were tested for hardness and the measurement was done according to BS EN ISO 6507-1-2018 standard with a load of 5 kgf. The variation in hardness is shown in Figure 7. Three layers from the top and bottom of the walls were not measured as they were sectioned out before grinding the sample surface.

The variation in hardness is shown in Figure 7. Three layers from the top and bottom of the walls were not measured as they were sectioned out before grinding the sample surface. As observed from the plot, the hardness varied between 140-160 HV and remained consistent along the deposition height, as reported in other thick wall depositions of low alloyed steels [13].Figure 7: Variation of hardness along the deposition height.1496Mechanical testing:Tensile samples extracted from the wall shown in Figure 2, were tested according tothe BS EN ISO 6892-1-2019 standard for two different material orientations.Figure 8: Yield and ultimate tensile strength values for W1 and W2 across two different orientations.The tensile properties are shown in Figure 8. Irrespective of the control methods (interlayer temperature or dwell time) and material orientation, W1 and W2 had similar values of the ultimate tensile strength of 480MPa and 470MPa, respectively.

Irrespective of the control methods (interlayer temperature or dwell time) and material orientation, W1 and W2 had similar values of the ultimate tensile strength of 480MPa and 470MPa, respectively. The yield strength values were 345 MPa and 332 MPa for W1 and W2, respectively. This corresponds well with previous observations by testing samples extracted from WA-DED manufactured low alloyed steel walls [11, 12].

This corresponds well with previous observations by testing samples extracted from WA-DED manufactured low alloyed steel walls [11, 12]. The lack of anisotropy in tensile properties can be due to the absence of preferred orientation (Figure 5) and the presence of equiaxed ferrite grains in the microstructure [13].Figure 9: Charpy impact test result for samples extracted from Walls 1 and 2 in two different orientations, notch parallel or notch perpendicular to build direction, tested across the transition temperature range.Charpy impact testing was conducted for standard-sized samples extracted from two different orientations, notch parallel or notch perpendicular to the build direction according to1497the BS EN ISO 148-1-2010 standards. Prior to the impact testing, the samples to be tested in temperatures below the room temperature where immersed in methyl alcohol bath for 4 minutes at the target temperature.

Prior to the impact testing, the samples to be tested in temperatures below the room temperature where immersed in methyl alcohol bath for 4 minutes at the target temperature. After attaining the test temperature, the samples were shifted to the testing chamber, where they were impact tested. Testing was done across a range of temperatures to determine the transition temperature and the temperature for different absorbed energies are given in Table 6.Table 6: Transition temperature values (°C) for the corresponding absorbed energy values for the impact samples tested.W1W2SampleHorizontal (Notch perpendicular)Vertical (Notch parallel)Horizontal (Notch perpendicular)Vertical (Notch Parallel)T27J(°C)83.870.398.392.6For the samples tested at room temperature, the samples tested from W1 and W2 had absorbed energy values of ~200J.

Comparing this result from literature on wire-arc DED depositions tested at room temperature, sub-sized specimens had lower impact energies [14] while regular sized samples had absorbed energy values of ~250J with observations of slight scatter in the toughness values [15, 16]. A wide scatter is observed for samples manufactured by the parallel deposition strategy [16], and it has been correlated with the microstructure being dependent on the phase formation. Rapid cooling after re-austenization led to Martensitic- Austenitic (MA) constituents which drastically influenced the scatter in impact toughness values, while slower cooling rate lead to fine pearlites formation enhancing the toughness values [16]. Overall, the impact toughness values are better for W2 on comparison with W1. The impact values demonstrate slight anisotropy, with notch perpendicular performing better than notch parallel samples for W1.

The impact values demonstrate slight anisotropy, with notch perpendicular performing better than notch parallel samples for W1. In comparison, W2 has lower transition temperature values that are consistent across testing orientation. The impact properties obtained in this study display better result owing to the slower cooling rate during deposition (from Table 4) that leads to fine equiaxed microstructure with no preferred orientation (Figure 5 and Figure 6).ConclusionThe microstructure and mechanical properties of WA-DED samples deposited using two interpass process control methods were characterised. On comparison, the two depositions have exhibited similar microstructural features, manifested by equiaxed ferrite grains without preferred grain orientation. The deposition made with interpass dwell time control displayed a slightly wider grain size variation than the deposition made with interpass temperature control.

The deposition made with interpass dwell time control displayed a slightly wider grain size variation than the deposition made with interpass temperature control. The mechanical properties (i.e., hardness, tensile strength and impact properties) of the two depositions were similar irrespective of the material orientation. However, slight anisotropy was observed in the Charpy impact transition temperature for both depositions, with lower transition temperature observed in the notch perpendicular direction. Furthermore, the deposition made with interpass dwell time control exhibited a lower overall (ie in both orientations) Charpy impact transition temperature. This suggest that using interpass dwell time control is as effective as interpass temperature control for low alloyed steel depositions through the WA-DED method.

This suggest that using interpass dwell time control is as effective as interpass temperature control for low alloyed steel depositions through the WA-DED method. However, this could be exclusive for thick wall depositions made through low alloyed steels which does not undergo complex phase transformations other than austenite to ferrite and carbide transition on reheating. Adopting dwell time method could be1498beneficial as it reduces the overall deposition time (~2 hours), thereby increasing process efficiency while eliminating the need for additional equipment for monitoring the surface temperature during the process.AcknowledgementsThis publication was made possible by the sponsorship and support of Lloyd’s Register Foundation. Lloyd’s Register Foundation helps to protect life and property by supporting engineering-related education, public engagement and the application of research.

Lloyd’s Register Foundation helps to protect life and property by supporting engineering-related education, public engagement and the application of research. The work was enabled through, and undertaken at, the National Structural Integrity Research Centre (NSIRC), a postgraduate engineering facility for industry-led research into structural integrity established and managed by TWI through a network of both national and international Universities. The work was also supported by Coventry University.Authors would like to acknowledge TWI Ltd and Armouers & Braisiers’ Gauntlet trust for awarding a travel grant that aided with the travel arrangements for participation at the conference.References[1] Rodrigues, T., Duarte, V., Miranda, R., Santos, T., & Oliveira, J. (2019). Current status and perspectives on wire and arc additive manufacturing (WAAM). Materials, 12(7), 1121. https://doi.org/10.3390/ma12071121[2] Ermakova, A., Mehmanparast, A., Ganguly, S., Razavi, J., & Berto, F.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperComparison of Rotating-Bending and Axial Fatigue Behaviors of LB-PBF 316LStainless SteelRakish Shrestha1,2, Jutima Simsiriwong3*, Nima Shamsaei1,21Department of Mechanical Engineering, Auburn University, Auburn, AL 2National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 3School of Engineering, University of North Florida, Jacksonville, FLCorresponding author: j.simsiriwong@unf.eduAbstractAdditive manufactured (AM) materials are prone to internal defects such as entrapped gas pores and lack of fusions along with having a rough surface. There are different types of fatigue tests that are used to characterize the effects of such defects on the structural integrity of AM parts.

There are different types of fatigue tests that are used to characterize the effects of such defects on the structural integrity of AM parts. The present study aims to investigate the effect of stress gradient on the fatigue behavior of 316L stainless steel (SS), fabricated using a laser beam powder bed fusion (LB-PBF) process. Axial fatigue tests are performed on as-built (non-machined) LB-PBF 316LSS round specimens with uniform gage section, while rotating bending fatigue tests are conducted on hourglass specimens (i.e. reduced gage section). Fatigue tests revealed that the specimens subjected to the axial loading exhibited lower fatigue resistance compared to the specimens failed under rotating bending test. Such differences in the fatigue life was attributed to the variation in the stress distribution resulting from different loading types and its effect on the fatigue crack propagation.

Such differences in the fatigue life was attributed to the variation in the stress distribution resulting from different loading types and its effect on the fatigue crack propagation. Fractography analysis conducted to determine the failure mechanism showed that all of the cracks initiated from the surface of the specimen irrespective to the loading conditions. Furthermore, fracture surface observed for LB-PBF 316L SS specimens resembled a typical fracture surface of notched specimens, which supports the fact that for the as-built specimens cracks initiates from the micro- notches as a result of layer wise fabrication in AM process.Keywords: Stress gradient, Fatigue, 316L stainless steel, Surface roughnessIntroductionLayer wise fabrication technique used in additive manufacturing (AM) process has provided design freedom for fabricating parts with complex geometries as well as complete assemblies.

As a result, AM technology is attracting interest from automotive and aerospace industries as a cost effective manufacturing process with an ability to fabricate lightweight parts to enhance fuel efficiency and reduce carbon foot print [1-3]. With an increased usage of AM techniques in various applications, it is necessary to establish standards to qualify and certify AM parts with an aim to ensure their reliability, especially under the realistic cyclic loading conditions [4].

Various standardization and federal agencies, such as American Society of Testing and Materials (ASTM) International, National Institute of Standards and Technology (NIST), Federal Aviation Administration (FAA), etc., are also actively involved in developing necessary guidelines515to qualify and certify AM parts with optimal mechanical properties under various loading conditions.Axial and rotating bending loadings are two of the most common loading types used for fatigue characterization of metallic materials. Furthermore, it has been recognized that the major differences in the fatigue behavior under axial and rotating bending tests arises from the differences in stress distribution experienced by the specimen. Normal stresses under axial loading is distributed uniformly throughout the gage volume of the specimen, while under rotating bending loading, maximum normal stress occurs at the surface and decreases linearly to zero at the center of the specimen [5, 6].

A correlation factor ranging between 0.75 and 0.9 has also been established to relate the fatigue limit of metallic parts subjected to axial and rotating bending loading [5]. In addition, this variation in stress distribution can affect the crack propagation behavior once the crack has been initiated [6]. Manson [6] also revealed that the nominal stress values used for rotating bending tests varies from the true stress due to the plastic deformation, which can have a significant effect on the fatigue life in low cycle fatigue regime.Sakai et al. [7] compared the effect of surface defects and sub-surface defects (i.e. an inclusion) on the fatigue behavior of wrought high carbon-chromium steel specimens subjected to axial and rotating bending loading. Specimens subjected to rotating bending loading exhibited higher fatigue lives when compared to the specimens subjected to axial loading at identical value of stress amplitude.

Specimens subjected to rotating bending loading exhibited higher fatigue lives when compared to the specimens subjected to axial loading at identical value of stress amplitude. Such result was attributed to the fact that under axial loading condition, entire volume of the specimen’s gage section experiences maximum stress, while under rotating bending loading, only the surface of the specimen experiences maximum stress. As a result, for the uniformly distributed defects, the number of critical defects becomes larger in the case of specimen under axial loading and consequently decreases the fatigue resistance. Furthermore, although higher fatigue lives were exhibited by the rotating bending specimens, the location of the defects (i.e.

Furthermore, although higher fatigue lives were exhibited by the rotating bending specimens, the location of the defects (i.e. surface or sub-surface) was reported to be more influential for rotating bending loading compared to the axial loading condition [7].The relationship between the fatigue behavior obtained from rotating bending and axial fatigue tests by understanding the effects of macroscopic and microscopic features for traditionally manufactured metallic materials are established. However, the effect of different types of defects inherent to additive manufacturing, such as surface roughness, lack of fusion (LoF), and gas entrapped pores [8, 9], on the type of loading is not yet understood. Furthermore, these defects in AM materials are not always uniformly distributed which could add more complexities to the fatigue analysis.

Furthermore, these defects in AM materials are not always uniformly distributed which could add more complexities to the fatigue analysis. Therefore, addressing this knowledge gap is necessary for understanding property-structure relationship and establishing the trustworthiness of fatigue and fracture critical AM parts in various industrial applications.In this study, the synergistic effects of surface roughness and loading types on the fatigue behavior of LB-PBF 316L SS are investigated. The material selected in this study, 316L SS is bio- compatible and exhibits excellent corrosion and wear resistance. Hence, 316L SS are currently being used in marine applications, nuclear power plants, and internal fixation implants in biomedical applications [10-12]. Following the introduction in this study, the material and experimental procedures are presented. The experimental results are then discussed with accompanied fractography analysis.

Following the introduction in this study, the material and experimental procedures are presented. The experimental results are then discussed with accompanied fractography analysis. Finally, important conclusions are drawn.516Material and Experimental Procedures316L SS powder produced by LPW Inc. with size distribution ranging from 15-45 μm was used to fabricate fatigue specimens utilizing a Renishaw AM 250, a laser beam powder bed fusion (LB-PBF) system, under argon purged environment. Manufacturer recommended process parameters (laser power of 200 W, hatch spacing of 0.11 mm, exposure time of 80 ms, and layer thickness of 50 μm) were used for the part fabrication. In this study, two sets of specimens, (I) cylindrical specimen with uniform gage section for axial fatigue test and (II) cylindrical specimen with reduced gage section (i.e. hourglass specimen) for rotating bending fatigue test, were designed according to ASTM Standard E-466 [13] and ISO 1143 [14], respectively.

hourglass specimen) for rotating bending fatigue test, were designed according to ASTM Standard E-466 [13] and ISO 1143 [14], respectively. The specimen dimensions are shown in Fig. 1(a) and 1(b) for axial and rotating bending fatigue tests, respectively. It should be noted that these specimens were designed to have similar cross section as shown in this figure. Additionally, the fatigue stress concentration factor for the rotating bending specimens was also calculated to be very small (Kf = 1.009). The specimens were manufactured in their net shaped geometry in vertical direction (i.e. perpendicular to the build plate). All fatigue specimens were further annealed at 1900 °F for one hour in vacuum followed by air cooling, while no post fabrication machining was conducted.I R38.10(a)I R38.10</)9.53 ~ </)4.76~(b)Figure 1. Specimen geometries and dimensions for (a) axial and (b) rotating bending fatigue tests.

Specimen geometries and dimensions for (a) axial and (b) rotating bending fatigue tests. All dimensions are in mm.To investigate the effect of stress gradient on the fatigue behavior of LB-PBF 316LSS in the presence of surface roughness, as-built specimens were subjected to axial and rotating bending fatigue loading. Rotating bending tests were conducted on a System Integrators RBF200 system under constant frequency, while MTS landmark servo-hydraulic system with 100 kN load cell was used to conduct uniaxial tests. All fatigue tests were conducted under fully-reversed (R(cid:305) = -1) force-controlled mode with stress varying between 250 – 375 MPa.

All fatigue tests were conducted under fully-reversed (R(cid:305) = -1) force-controlled mode with stress varying between 250 – 375 MPa. Detail fractography analysis on the specimens failed under both axial and rotating bending tests were further analyzed using Keyence VHX 6000 digital microscope to determine the factors responsible for crack initiation as well as failure mechanism arising from variation in fatigue loading.517Experimental Results and DiscussionsStress amplitude versus fatigue life data obtained for LB-PBF 316L SS specimens subjected to axial and rotating bending cyclic loadings are presented in Fig. 2. A significant effect of loading type on the fatigue behavior of LB-PBF 316L SS was observed, as specimens generally exhibited better fatigue resistance under rotating bending loading when compared to the axial loading condition.

At higher stress amplitudes, the difference in fatigue life was seen to be more than an order of magnitude, while at lower stress amplitudes, the difference was within approximately a factor of three.5000 Rotating Bending450♦ Axial400;- 0. ~ 350 a, "C -~ c.. ~ 250 "' "' ~ 200 in 100300+-+-+++t+t. 1 1i [ I [1E+031E+041E+051E+061E+07Reversals to Failure, 2N1Figure 2. Stress amplitude versus fatigue life data obtained from rotating bending and axial loading under force-controlled mode.Schematic representing the normal stress distribution on cylindrical specimens under rotating bending and axial loading conditions is shown in Fig. 3(a) and 3(b), respectively. Under rotating bending loading, the stresses are highest at the surface of the specimen, and gradually decrease and reach zero at its neutral axis, while under axial loading, the stress distribution remains constant throughout the entire gage section of the specimen.

This stress gradient effect can influence the crack initiation and crack growth behavior [7]. In high cycle fatigue regime, the majority of the fatigue life is dominated by crack initiation stage. Since the cracks mainly initiate from the surface of AM specimens in the presence of surface roughness [15, 16], this may have resulted in smaller variation in the fatigue lives in lower stress amplitude levels as indicated in Fig. 2. It is also important to mention that, based on the specimen geometry, axial specimens with uniform gage section have higher gage volume when compared to the rotating bending specimens with reduced gage section. As a result, axial specimens have higher probability of having larger amount of crack initiating defects as compared to the rotating bending specimens, which consequently can also reduce the fatigue life.In low cycle fatigue regime (i.e. higher stress amplitude levels), the majority of fatigue lives is spent in crack growth regime.

higher stress amplitude levels), the majority of fatigue lives is spent in crack growth regime. As a result, once the crack initiates in the specimen subjected high stress amplitudes, the crack starts to propagate inwards in the direction towards the center of the specimen. Since there is less stress within the internal region of the rotating bending specimen’s518cross section as compared to its surface as shown in Fig. 3(b), crack will grow in a slower rate. On the other hand, since the stress is uniform throughout the gage section of the specimen under axial loading, crack will grow at similar rate. This effect of stress gradient may be responsible for the larger variation in the fatigue lives obtained from rotating bending and axial loading conditions under identical stress levels at both high and low cycle fatigue regimes.(a) (b) Figure 3.

Schematic showing distribution of stress under (a) rotating bending and (b) axial loading.To further investigate the failure mechanisms of LB-PBF 316L SS under different loading conditions, fractography analysis was conducted on selected specimens subjected to rotating bending and axial loading at identical stress levels. Fracture surfaces obtained for specimens subjected to the stress amplitude of 350 MPa under rotating bending and axial fatigue loading are shown in Fig. 4(a), and 4(b), respectively. Irrespective to the loading condition, cracks were always observed to initiate from multiple location as a result of surface roughness in all specimens. Upon further magnification, fatigue cracks were seen to initiate from several micro notches located at the surface of the specimen as indicated by the red arrows in the zoomed in images. Furthermore, the fracture surfaces seen for LB-PBF 316L SS also resembled a typical failure mechanism observed in notched specimens shown in [5].

Furthermore, the fracture surfaces seen for LB-PBF 316L SS also resembled a typical failure mechanism observed in notched specimens shown in [5]. Therefore, in as-built surface condition, micro notches resulting from the layer wise fabrication process of AM technique may be the major factor driving the fatigue failure in AM parts.(a)519(b)Figure 4. Fracture surfaces showing differences in failure mechanism in as-built specimens (i.e. non-machined) subjected to (a) rotating bending and (b) axial loadings.ConclusionsIn this study, the effect of stress distribution within the volume of the specimens’ gage section on the fatigue lives of LB-PBF 316L SS in as built (non-machined) surface condition was investigated. Based on the results obtained and discussions presented, the following conclusions were drawn:1. Some influence of stress gradient was observed on the fatigue behavior of LB-PBF 316L SS.

Based on the results obtained and discussions presented, the following conclusions were drawn:1. Some influence of stress gradient was observed on the fatigue behavior of LB-PBF 316L SS. In general, higher fatigue resistance was observed for specimens subjected rotating bending loading compared to those subjected to axial loading.2. Larger difference in the fatigue lives resulting from the variation of fatigue loading was seen at low cycle fatigue regime as compared to high cycle fatigue regime. This was attributed to the influence of stress gradient on the crack growth behavior, which governs the majority of fatigue lives at low cycle fatigue regime.3. Fractography analysis revealed that the cracks always initiated from the micro-notches present at the surface of the specimen irrespective to loading condition. This result was also supported by the fracture surface that resembled those of notched specimen made of wrought counterpart.

This result was also supported by the fracture surface that resembled those of notched specimen made of wrought counterpart. AcknowledgementsThis material is based upon work supported by the National Science Foundation (NSF) under Grant No. 1657195. Authors would also like to thank Mr. Ross Wykoff and Mr. Daniel Hoover, undergraduate students at the University of North Florida, for assisting rotating-bending fatigue tests. In addition, Authors would like to acknowledge Braddock Metallurgical, Jacksonville, FL for performing heat treatments for all specimens used in this study.References520[1] K.V. Wong and A. Hernandez, A review of additive manufacturing, ISRN MechanicalEngineering, (2012) doi:10.5402/2012/208760.[2] R. Huang, M. Riddle, D. Graziano, J. Warren, S. Das, S. Nimbalkar, J. Cresko, and E. Masanet, Energy and emissions saving potential of additive manufacturing: the case of lightweight aircraft components, Journal of Cleaner Production, 135 (2016) 1559-1570.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceA Comparison of Stress Corrosion Cracking Susceptibility in Additively-Manufactured and Wrought Materials for Aerospace and Biomedical ApplicationsMichael Roach1, R. Scott Williamson1, Jonathan Pegues2,3, Nima Shamsaei2,3,*1Department of Biomedical Materials Science, University of Mississippi Medical Center, Jackson, MS 39216 2Department of Mechanical Engineering, Auburn University, Auburn, AL 368493National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849Corresponding author: shamsaei@auburn.eduAbstractAdditive manufacturing (AM) is becoming an increasingly popular method in both aerospace and biomedical industries. Titanium alloys are increasingly common in additive manufactured applications due to their excellent strength to weight ratio and biocompatibility.

Titanium alloys are increasingly common in additive manufactured applications due to their excellent strength to weight ratio and biocompatibility. Traditional wrought Ti-6Al-4V alloys show little sensitivity to stress corrosion cracking (SCC) when subjected to in-vitro conditions. In AM applications the alloy powder is often sifted and reused multiple times which often results in a degradation of the powder’s shape. Recent studies have also shown that the oxygen content of additive powders increases with repeated powder use which may increase the susceptibility of the resulting parts to SCC. This research compares the microstructural characteristics and tensile SCC behavior of AM Ti-6Al-4V parts fabricated from new and recycled powder in distilled H2O, salt water, and Ringers solution.

This research compares the microstructural characteristics and tensile SCC behavior of AM Ti-6Al-4V parts fabricated from new and recycled powder in distilled H2O, salt water, and Ringers solution. Additionally, the effect of surface finish is investigated for each microstructure comparing the as-built surfaces to machined and polished surfaces.Keywords: Stress corrosion cracking, environmentally assisted cracking, titanium alloy, tensile propertiesIntroductionAs additively manufacturing becomes a more viable option for fabrication of structural components in a wide range of biomedical and aerospace applications there is a critical need to understand the failure mechanisms related to these specific applications. As a result, numerous studies have recently been published comparing mechanical behavior of additively manufactured (AM) titanium alloy parts to their wrought material counterparts [1]–[13].

As a result, numerous studies have recently been published comparing mechanical behavior of additively manufactured (AM) titanium alloy parts to their wrought material counterparts [1]–[13]. Despite this recent surge1410of research the effect of environmentally assisted cracking (EAC) and its effect of stress corrosion cracking (SCC) has not been well represented for additive manufactured (AM) materials.Titanium alloys such as Ti-6Al-4V are common in both biomedical and aerospace industries due to their high strength to weight ratio, biocompatibility, and excellent corrosion resistance. In both of these industries parts are often subjected to harsh environments that may lead to an increased susceptibility to SCC [14], [15].

In both of these industries parts are often subjected to harsh environments that may lead to an increased susceptibility to SCC [14], [15]. While the EAC failure mechanisms are well reported for wrought Ti-6Al-4V and believed to be of relatively low risk, however, the unique microstructure, equilibrium phases, and high surface roughness associated with AM parts could lead to an increased risk of SCC compared to its wrought counterpart. Recent studies have also shown that the oxygen content of additive powders increases with repeated powder use which may increase the susceptibility of the resulting parts to SCC [16], [17].The purpose of the present research was to compare the ASTM G129 slow strain rate (SSR) SCC susceptibility of Ti-6Al-4V samples fabricated using laser powder bed fusion (LPBF) technology, to those fabricated using traditional machined wrought samples from the same alloy.

The ductility of AM Ti-6Al-4V samples were compared to their wrought counterparts in distilled water, Ringer’s, and 3.5% NaCl electrolytes. Additionally, the SCC response of titanium AM samples fabricated from the virgin and reconditioned powders were compared for both as-built and machined conditions.Experimental Set-upTest coupons were designed accordance to ASTM G129 [18] with the dimensions shown in Figure 1. Grip ends were threaded in order to submerge the test coupons in the desired solution during testing. Two solutions, Ringer’s and 3.5% NaCl, were selected to test for stress corrosion cracking susceptibility compared to distilled H2O as the control solution. Ringer’s solution is often used in testing to simulate physiological conditions. Additionally, a 3.5% NaCl solution is often used in testing to simulate a seawater environment.

Ringer’s solution is often used in testing to simulate physiological conditions. Additionally, a 3.5% NaCl solution is often used in testing to simulate a seawater environment. For the Ringer’s solution, a test temperature of 37°C was chosen as a representative in-vivo condition for surgical implants while room temperature (23°C) was selected for testing in the 3.5% NaCl solution. For the control solution, a test temperature of 37°C was used.1411Figure 1: Specimen geometry used for slow strain rate testing.Slow strain rate tensile tests were performed on an MTS load frame with a 100 kN capacity. A displacement stroke rate of 2.5x10-5 mm/s was chosen which results in an approximate strain rate of 1x10-6 s-1. Test specimens were completely submerged in the test solution and the temperature was monitored and controlled using a Digi-Sense temperature controller. Figure 2 gives a representation of the test setup used for this study.

Figure 2 gives a representation of the test setup used for this study. All samples were tested until failure, after which, a measuring microscope was used to determine the elongation to failure and reduction of area (RoA). Fracture surfaces were then examined with both a Keyence digital microscope and a Zeiss scanning electron microscope (SEM). Ratios between the control environment and the harsh environment were calculated by dividing the mean values of samples in harsh environments by the mean values of the samples in the control (distilled H2O). ASTM G129 [18] states that ratios that are less than unity are considered to be susceptible to SCC.

ASTM G129 [18] states that ratios that are less than unity are considered to be susceptible to SCC. For this study a more strict definition of ratios falling below 0.9, which has been used in previous studies [14], is considered to be susceptible to SCC.1412Figure 2: Test set up showing specimen submerged in solution during testing.ResultsThe elongation to failure, reduction of area, and time to failure were averaged for each condition and are listed in Table 1. The elongation and reduction of area are significantly lower than the wrought conditions which is consistent with literature on L-PBF Ti-6Al-4V [5], [13], [19]. The low ductility of AM L-PBF Ti-6Al-4V is attributed to the fine acicular martensitic microstructure as a result of the high cooling rate during fabrication.

The low ductility of AM L-PBF Ti-6Al-4V is attributed to the fine acicular martensitic microstructure as a result of the high cooling rate during fabrication. Rapid solidification in the fabrication process does not allow the diffusive a transformation to occur, thus, the β phase transforms to the α’ phase through the non-diffusive martensitic transformation [20]. Additionally, the elongation to failure for all conditions reported in this study is much lower than the required 10% elongation to failure required by ASTM F2924 [21] for AM Ti-6Al-4V.Table 1: Slow strain rate results for each condition.ConditionPowderSolutionTemp.

(°C)Elongation (%)RoA (%)As-BuiltNewH2O376.2728.2As-BuiltReusedH2O376.1025.21413Time (Hrs)41.2940.15As-BuiltNewRingers375.5823.3As-BuiltReusedRingers375.3125.2As-BuiltNew3.5% NaCl235.3021.7As-BuiltReused3.5% NaCl234.9723.9MachinedNewH2O375.9231.9MachinedNewRingers375.8334.0Wrought [1]NAH2O3721.853.7Wrought [1]NARingers3720.553.3A visual representation of the slow strain rate tensile properties for each condition is shown using a box plot in Figure 3. The dotted lines on each box plot represents the lower range of the control solution with the black line representing the machined control condition and the red line representing the as-built control condition. The elongation to failure was mostly similar for all conditions and solutions with only the 3.5% NaCl solution showing significant decreases.

The elongation to failure was mostly similar for all conditions and solutions with only the 3.5% NaCl solution showing significant decreases. There was a slight decrease in elongation to failure for samples in Ringer’s, however, the difference is not significant. A similar trend was observed for the reduction of area (RoA) with Ringer’s and 3.5% NaCl showing slightly lower RoA. Interestingly, while no difference was observed between as-built and machined specimens in elongation to failure, as-built specimens did show consistent decreases in RoA compared to machined specimens.

Finally, there was no significant differences in the time to failure for all conditions.141440.7939.9338.3338.1639.1241.01NANAFigure 3: Box plots showing the spread in data for all conditions.Evaluation of environmentally assisted cracking for slow strain rate testing, according to ASTM G129 [18], is carried out by comparing the SSR ratios to the control environment with ratios reaching unity being considered not susceptible to SCC. Decreasing SSR ratios are considered to show increasing susceptibility to SCC. In general, for biomedical applications, ratios that are lower than 0.9 are considered to show some susceptibility to SCC [14]. Table 2 lists the elongation ratio (RE), the reduction of area ratio (RRA), and time to failure ratio (RTTF) calculated1415for each condition.

Table 2 lists the elongation ratio (RE), the reduction of area ratio (RRA), and time to failure ratio (RTTF) calculated1415for each condition. The control for all machined specimens is the machined new powder in distilled H20 while the control for all as-built conditions was the as-built new powder in distilled H20.Table 2: Calculated ratios for each condition, * represents control for surface condition.ConditionPowderSolutionTemp.

(°C)RERRARTTFMachinedNewH2O371.001.001.00MachinedNewRingers370.981.071.05MachinedNew3.5% NaCl230.870.930.97As-BuiltNewH2O371.001.001.00As-BuiltReusedH2O370.970.890.97As-BuiltNewRingers370.890.830.99As-BuiltReusedRingers370.850.900.97As-BuiltNew3.5% NaCl230.840.770.93As-BuiltReused3.5% NaCl230.790.850.92Comparing the ratios for each condition given in Table 2, there are several areas that indicate susceptibility to SCC for L-PBF Ti-6Al-4V. For the machined conditions there is a decrease in the elongation to failure in the 3.5% NaCl solution, however, the reduction of area is only slightly reduced. The most notable observation is the significant decrease in ductility for all as-built conditions.

The most notable observation is the significant decrease in ductility for all as-built conditions. While the elongation to failure ratio was not significantly reduced for distilled H2O fabricated from both new and reused powder, all other as-built geometries showed significant (< 0.9) decreases in RE. Additionally, the RRA was also considered to be significant (< 0.9) for all as-built surfaces outside of the control condition. This indicates that the high roughness associated with the as-built geometries could potentially increase the susceptibility of EAC for AM Ti-6Al-4V. Additionally, all 3.5% NaCl environments showed some loss in ductility compared to the control environment with significantly low RRA (≤ 0.75) for the as-built conditions. The reduced ductility for the Ringer’s and 3.5% NaCl conditions suggests that AM Ti- 6Al-4V parts could potentially show an increased risk of EAC in harsh environments in the presence of high surface roughness.

Additionally, another observation than can be made from Table 2 is the further decrease in ductility for all 3.5% NaCl solutions. More tests for specimens in 3.5% NaCl at 37° C and distilled H2O at 23°C are necessary to ensure the difference in testing temperature is not playing some role in the SCC behavior.Fractography analysis revealed a change in crack behavior between the control and 3.5% NaCl solutions while the Ringer’s solution was mostly similar to the control. Representative fracture surfaces for each condition are shown in Figure 4. For the control condition in Fig. 4(a) the crack propagation appears to be deflect along prior β grains as indicated by the deep valleys. Closer inspection of these valleys to the right of Fig. 4(a) reveals that these appear to be ends of1416prior β grains being pulled completely from the surface. This type of crack propagation results in multiple crack deflections leading to a tortuous fracture surface.

This type of crack propagation results in multiple crack deflections leading to a tortuous fracture surface. Similarly, the fracture surface of the representative Ringer’s solution in Fig. 4(b) shows multiple crack deflections along the prior β grains resulting in a tortuous fracture surface. Conversely, the 3.5% NaCl fracture surface shows a less tortuous crack path which is attributed to very little failure along prior β grains. This type of crack propagation consists of less crack deflection ultimately leading to the lower ductility in the 3.5% NaCl specimens.0Figure 4: Fracture surfaces of the as-built condition in a) distilled H2O, b) Ringers, and c) 3.5% NaCl.1417ConclusionsThe results presented in this study indicate that additive manufactured Ti-6Al-4V may be susceptible to stress corrosion cracking.

A decrease in elongation to failure was observed for a solution of machined 3.5% NaCl while the machined Ringer’s solution showed similar elongation to failure as the control in the machined condition. However, all samples in the as-built condition showed significant reduction of area ratios (< 0.9) compared to the as-built control. The reduction in elongation to failure in the as-built conditions was observable but less severe than the reduction of area. This significant loss in ductility for the as-built specimens in Ringer’s and 3.5% NaCl compared to distilled H2O in the as-built condition indicate that the presence of the rough surface may increase the susceptibility of AM Ti-6Al-4V to SCC. These results, while preliminary, do indicate that stress corrosion cracking could be of concern for AM Ti-6Al-4V in the presence of harsh environments. Further tests are in progress to determine the effect of testing temperature in relation to the test solution.

Further tests are in progress to determine the effect of testing temperature in relation to the test solution. Additionally, further in-depth studies are needed to fully assess the effect of harsh environments on AM parts.AcknowledgementsPartial funding for this work was provided by the National Science Foundation under Grant No. 1657195. Supports from the Naval Air Systems Command (NAVAIR) is also greatly appreciated.References[1] B. E. Carroll, T. A. Palmer, and A. M. Beese, “Anisotropic tensile behavior of Ti–6Al–4V components fabricated with directed energy deposition additive manufacturing,” Acta Mater., vol. 87, no. Supplement C, pp. 309–320, Apr. 2015.[2] A. J. Sterling, B. Torries, N. Shamsaei, S. M. Thompson, and D. W. Seely, “Fatigue behavior and failure mechanisms of direct laser deposited Ti–6Al–4V,” Mater. Sci. Eng. A, vol. 655, pp. 100–112, Feb. 2016.[3] K. S. Chan, M. Koike, R. L. Mason, and T.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperComparison of Surface Texture from Various Surface Morphology Techniques for Evaluating As-Built Ti6Al4V Laser Powder Bed FusionAlex De La Cruz1, Francisco Medina1, Edel Arrieta1, Luke Weston2, Mark Benedict2, Thao Gibson31 University of Texas El Paso, El Paso, TX 2 Air Force Research Laboratory, WPAFB, OH 3 University of Dayton Research Institute, Dayton, OHAbstractAdditive manufacturing (AM) is capable of creating unique and complex geometries thatconventional methods cannot achieve. The applications for AM have been rapidly increasing across a variety of sectors, particularly for biomedical and aerospace components, the relatively low production volumes and high demand for customizability in both sectors are especially amiable to AM.

However, without post-processing, AM components contain a variety of flaws, such as surface roughness and porosity, that can partially be mitigated by process parameters like scan speed and laser power. Surface roughness is a flaw present for every as-built AM surface that serves as an array of sites for every mode of material failure to occur. Common surface roughness measurements involve the use of optical and contact stylus profilometry. However, x- ray Computed Tomography (xCT) is already the most widely used method of analyzing AM parts for porosity, inclusions, and various other flaws. In terms of resolution, xCT should be fully capable of analyzing surface roughness and is the only method of the three investigated that can inspect interior geometries. Therefore, evaluating xCT as a fully inclusive analysis method for AM parts is advantageous. In this study, we compared three surface characterization technologies, xCT, optical profilometry, and contact stylus profilometry.

In this study, we compared three surface characterization technologies, xCT, optical profilometry, and contact stylus profilometry. The comparison of these technologies is being done on as-built Laser Powder Bed Fusion (L-BPF) Ti6AI4V four- point bending fatigue samples. Further understanding the difference among each of the technologies will aid ongoing research on developing a standard for xCT surface characterization while also providing more knowledge and insight into each technique and what can be expected. Each of the samples was produced by varying scanning speed and laser power, resulting in different surface textures.

Each of the samples was produced by varying scanning speed and laser power, resulting in different surface textures. Preliminary results show deviations of Sa _%, Sz _%, Sv _%, and Sku _% between the xCT and optical microscopy methods are comparable between these two methods.1.IntroductionLaser Powder Bed Fusion (L-PBF) is an Additive Manufacturing (AM) method thatbuilds a metal component by melting metal powder particles layer by layer via a laser source or electron beam until the desired geometry is completed [1], [2]. This method brings liberty to the1756user by manufacturing complex and unique geometries that traditional manufacturing methods cannot obtain. For instance, AM processes are additive in the sense that they add material to create the part design rather than removing material as is the case with subtractive processes such as milling or turning [3].

This emerging technology could be used to produce tools and components at a very high rate for a wide range of industries, including biomedical, aerospace, and automotive [2]. Such materials that are prominent in the industry for their high performance are titanium (Ti6Al4V), 17-4 PH stainless steel, cobalt chrome, and Inconel 625, which have been shown to be suitable materials for AM production [4]. The overall uses of AM components depend on several factors such as their effectiveness, durability, and overall performance. These influences highly depend on their general surface texture and the surface characterization of their defects, which have the potential to alter the performance efficiency and quality of the component by reducing their durability as well as life expectancy [5].

Additionally, the as-built surface texture is a reflection of the L-PBF machine configurations, such as the powder particle size, layer thickness, beam power, incident angle, and build orientation, all of which are determined before the building process [6]. As-built surfaces of AM-produced components tend to have higher surface roughness, which results in more valleys and peaks that act as stress concentration zones, resulting in a less efficient part [5].

The industry is still in need of a standard that can be utilized to consistently quantify and be able to anticipate the performance according to the L-PBF printing parameters as well as surface texture characterization, notwithstanding the difficulty of measuring the surface texture of AM parts [6].As the surface texture is an essential and valuable measure, a study was performed wherea comparison of different and most commonly used optical measuring technologies such as Confocal Microscopy (CM), Coherence Scanning Interferometry (CSI), Focus Variation Microscopy (FVM), Computed X-ray tomography (XCT) was done to compare their discrepancies [4]. Optical technologies offer a non-destructive surface measuring method of the surface area where post-processing of this data is needed to determine surface roughness measurement Sa.

Optical technologies offer a non-destructive surface measuring method of the surface area where post-processing of this data is needed to determine surface roughness measurement Sa. In contrast, contact methods such as a profilometer use a probe tip that is placed on the surface of the sample and records a line profile that will measure the sample's roughness measurement, known as Ra [7]. Furthermore, the contact stylus has been widely used for the measurement of roughness and is the main reference instrument where it provides traceable measurements of form and surface texture [6]. CM offers high lateral and axial resolution, which are dependent on the objective magnification and numerical aperture [7]. A CM uses a pinhole disk in which the light from the LED source is focused through the pinhole disk and the objective lens onto the sample surface, which will reflect the light that is being projected [8].

The reflected light will be filtered by the pinhole disk, so only in-focus light reaches the camera sensor [8]. CSI, also known as white light scanning interferometry, is a technique based on far-field optical microscopy that uses white light interference fringes as a probe scanned over the depth of the sample in order to measure the surface roughness [9]. FVM utilizes an optic with a limited depth of field combined with a vertical scanning process where a series of images is1757recorded as the optic is moved vertically along the optical axis, resulting in a vertical image stack that later will be condensed and stitched together using a software algorithm to create a 3D image [10]. Moreover, the non-destructive method xCT has gained significant attention by allowing the user to inspect their manufactured AM parts' geometry internally and externally without the need to cut or destroy it [2].

However, some notable factors that affect the reconstruction of the xCT data and measurement are resolution, sample material, and machine features [2]. In addition, another factor that impacts the resolution, as well as the ability to measure the surface effectively, is the magnification and voxel size, which highly contribute to the final resolution of the surface [2]. A method that is used in xCT data analysis is surface determination from the gray value distribution, which allows for the separation of the material from the background using the gray histogram [2]. xCT and the implementation of this technology in the industry are rapidly growing industries in the use of this technology for the characterization of parts and other applications are the manufacturing, electrical, and food industries.As previously stated, Ra and Sa are two parameters that quantify the surface roughness of a component.

However, a clear distinction between both of these measuring parameters is that Sa is a surface area measurement, as opposed to Ra, which is a line profile measurement [2]. Sa, the areal parameter presents more insight into irregular surfaces of AM parts as it takes into account the whole analyzed areal topography not only a line profile [2]. AM manufacturing companies characterize their surfaces using Ra and Rz measurement methods, and little research shows the use of areal measurements and other parameters either with tactile or optical instruments [3]. From Triantaphyllou's study, it was concluded that Sa and Sq measurements were found to be suitable areal measurement parameters for AM surfaces when compared between two AM technologies, EBM and SLM. In this study, we analyzed the surface roughness Ra and Sa values using three different technologies: xCT, Keyence VR 5200, and SURFTEST SJ-210 Profilometer.

In this study, we analyzed the surface roughness Ra and Sa values using three different technologies: xCT, Keyence VR 5200, and SURFTEST SJ-210 Profilometer. The measurements were taken after samples were manufactured via an L-PBF machine at varying printing parameters. The chosen samples were later fatigue tested. These technologies are distinctive; however, this research will assist and provide more insight into the use of xCT to determine the surface roughness of AM components in anticipation of progressing and implementing this technology and standardizing this method. The comparison of three different surface measuring technologies is to provide a thorough comparison and quantify the accuracy of measuring surface roughness across different platforms of commonly used technologies.

In addition, this study will provide more insight when measuring surface roughness from AM-produced samples not on what technology is the best form of measurement but provide data and results on what technology would be best applicable to the application in hand.2.Materials and Methods2.1 Manufacturing of components1758Samples that were used in this study were manufactured via the L-PBF EOS M-290machine with dimensions of 5 mm x 5 mm x 70 mm. The primary material that was used for all samples was Ti6AI4V; however, the material density was not taken into account upon manufacturing to ensure xCT measurements were precise, as material density does have an effect on the X-ray absorption when conducting xCT measurements [2].

The samples were manufactured at varying printing parameters by altering the speed and power of the L-PBF machine laser source, which resulted in regions known as keyhole, process window, and lack of fusion, the power and speed specifications are illustrated on Fig. 1.Parameter Keyhole Process Window Lack of Fusion EOS NOM EOS NOM IMPower (W) 370 370 370 280 280Speed (mm/s) 800 1400 2000 1200 1200Fig 1. Printing parameter power and speed specifications per regimeKeyhole regions are called such due to the deeply penetrating heat-affected zone due tolow speed but high-power input, which resembles a keyhole in cross-section. Lack of fusion regions are regions with insufficient heat transfer due to high speed and low power which results in poor melted and fused together layers, this region tends to have rough edges, pores, and trapped powder. Process window regions are regions where mechanically desirable microstructures are produced.

Process window regions are regions where mechanically desirable microstructures are produced. The latter case is only possible in the “window” of process parameters that both create fully dense metal by fully and uniformly melting the metal without overheating it. Process window parameters vary between alloy systems and AM machines and are part of the reason AM is difficult to qualify in actual applications. In addition, we printed samples with EOS standard printing parameters, which resulted in two additional parameters being formed that are also within the process window regime. In this study, the two additional parameters from EOS standards will be labeled as EOS nominal (EOS NOM) and EOS nominal improved (EOS NOM IM). A total of five parameters were used and modifying and altering the power and speed of the laser will allow us to achieve different surface finishes on the as-built surfaces of the components [7], [11].

Furthermore, stress leaving was a post-process applied to each print to reduce the effect of internal thermal stresses after the manufacturing process. The stress-relieving temperature was set at 600ºC for 120 minutes with a heating and cooling rate of 5ºC/min.2.1 X-Ray Computed Tomography1759The xCT measurements were conducted using a ZEISS Xradia 620, where two voxelresolutions were used per sample, one at a 3.5 µm voxel resolution with a 0.4x magnification and the other at a 0.7 µm voxel resolution at a 4x magnification. The high voxel resolution (0.7 µm) was taken at an area of 1 mm x 1 mm where the larger voxel resolution (3.5 µm) was capable of capturing the samples four faces. For both of the voxel resolutions, the power and voltage remained consistent at 140 kV and 21 W.

For both of the voxel resolutions, the power and voltage remained consistent at 140 kV and 21 W. The 3.5 µm voxel resolution xCT machine parameters were set to the following: sample field of view of 7.4 mm2, detector distance to sample of 210 mm, source distance to a sample of 25 mm, exposure time of 4.5 sec per projection, and the source filter set at HE1. While the 0.7 µm voxel resolution samples were scanned with a sample field of view of 1.4 mm2, a detector distance to the sample of 77 mm, a source distance to the sample of 20 mm, an exposure time of 20 seconds per projection, and the source filter set to HE3. The source filters were chosen according to ZEISS guidelines and were determined based on the response to the transmission of the sample at 140 kV/21W.

The source filters were chosen according to ZEISS guidelines and were determined based on the response to the transmission of the sample at 140 kV/21W. The functions of the filter are used primarily for improving the reconstruction of the image quality by removing low-energy X- rays that passed through the sample [13].Upon completion of xCT, all datasets of each sample were analyzed using VGSTUDIO Max software from Volume Graphics. The procedure taken upon importing dataset files in VG- Studio Max was followed by altering the grayscale histogram and separating our material component from the background and excessive noise. We added a clipping-box tool, where the dataset was clipped to a designated area avoiding any excess material bleed causing inaccurate results when conducting surface analysis measurements. The clipped box was then converted into a Region of Interest (ROI), where a new volume was created by extracting the ROI.

The clipped box was then converted into a Region of Interest (ROI), where a new volume was created by extracting the ROI. An advanced surface-determination tool was incorporated into the model to properly separate the material and excess noise that was left behind due to the grayscale histogram, it has been also shown that the “integrated surface determination” setting shows a much better separation of material and background noise. [AR 4.13 on how the integrated surface determination is better]. We then converted the extracted modified surface to an STL file using high simplification to reduce the amount of computing power. The STL file was then imported into the Omnisurf surface analysis software from Digital Metrology. This method was useful for both voxel sizes of 0.7 µm and 3.5 µm; however, the 3.5 µm voxel resolution dataset contained all four sides of the component whereas the 0.7 µm voxels were taken at an area of 1 mm x 1 mm.

We followed the same procedure previously described and created an STL surface mesh for the 3.5 µm voxel resolution samples on the loading and supporting pin sides for the fatigue testing.The STL meshes were then imported into Omnisruf software which has the capability to analyze xCT surface meshes and output surface roughness (Sa) and line profile roughness (Ra) values. Upon adding the surface mesh, we re-cropped the STL mesh to a 750 µm x 750 µm area, this was done to ensure comparability with high voxel resolution scans which were at a 1 mm x 1 mm area with low voxel resolution as well as with Keyence measurements. After the cropping is1760done, we then ensure to use a bi-cubic fill which results in filling in any missing data the STL might be missing due to the exportation of the STL mesh.

After the cropping is1760done, we then ensure to use a bi-cubic fill which results in filling in any missing data the STL might be missing due to the exportation of the STL mesh. Gaussian S and L filters were also applied following ISO standard 25178-3 where an S-filter of 0.0025 mm and an L-filter of 2.5 mm resulted in a bandwidth ratio of 1000:1.2.3 ProfilometerA Surftest SJ-210 Mitutoyo surface roughness measuring tester was used in the analysisof each sample where the roughness value (Ra) was measured.

The specifics of the SJ-210 profilometer are as follows: a tip diameter of 5 µm, a cut-off filter of 0.8 mm, a transversal speed of 0.5 mm/s, with a total of 5x sampling lengths following ISO 4287 (1997) standard which was selected within the device's software.2.4 KeyenceThe Keyence VR 5200 was used to measure both Sa and Ra values from each sample,where the magnification were taken, at 80x, and the measurements were taken on the tension and compression sides, on which the supporting and loading pins made contact during the four-point bend test, respectively. The measurements were taken at three different locations in each parameter sample, and the average values are reported. A key factor when using the Keyence is the S-filter needs to be determined. The S-filter is an aerial gaussian filter used to create cutoffs for sampling distance and sphere radius according to ISO 25178-3 [12].

The S-filter is an aerial gaussian filter used to create cutoffs for sampling distance and sphere radius according to ISO 25178-3 [12]. These cutoffs were important in surface roughness analysis because the L-PBF samples were as-built surfaces, and some un-melted particles were observed on the surface visually as well as optically. Thus, a 50 µm S-filter was selected to measure only the roughness of the sample surface, rather than particles stuck to the sample.

Thus, a 50 µm S-filter was selected to measure only the roughness of the sample surface, rather than particles stuck to the sample. The Keyence was used for both Sa and Ra measurements for all samples where the Sa was taken in a selected area of 750 µm x 750 µm and the profile measurement was taken perpendicular to the build direction with a distance of 750 µm.3.Results1761Fig 2: Sa Comparison on 3.5 µm voxel xCT vs KeyenceFig 3: Sv Comparison on 3.5 µm voxel xCT vs KeyenceFig 4: Sz Comparison on 3.5 voxel xCT vs KeyenceFig 5: Sku Comparison on 3.5 voxel xCT vs KeyenceFig 6: Graphical representation of % different among xCT vs Keyence Surface roughness values (Sa, Sz, Sv, Sku)1762Fig 7: Percentage difference between 3.5 µm voxel xCT vs Keyence per each printed parameter set.4.DiscussionAs previously mentioned, the study was done to compare xCT, optical microscopy and profilometer.

While the data only show xCT at a 3.5 µm voxel resolution and Keyence optical microscopy comparison, current and ongoing work is being done to further compare all methods. Further investigation of the study will allow us to aid current research to progress the standardization of xCT method for external and internal measurements. In addition, provide more insight on each technology and their differences when evaluating surface roughness. Moreover, preliminary results show comparable data from xCT 3.5 µm voxel resolution with Keyence optical microscopy.Fig 2 – Fig 5 is a graphical comparison of parameters, Sa, Sv, Sz, and Sku. Where Sa is the difference in height of each point compared to the arithmetical mean of the surface, Sv is the absolute value of the height of the largest pit, Sz is the sum of the largest peak height value and the largest pit depth, Sku is the sharpness of the roughness also known as Kurtosis.

These evaluated parameters are measured within the respected surface area in which is being analyzed which in this case was a 750µm x 750 µm area. Additionally, Fig 6 illustrates the difference in percentage between the two measuring methods. Where Sv shows the highest differences up to 104.26% for the Keyhole parameter. A current assumption for such a significant difference is the Keyence uses an LED source where it can penetrate in a larger depth to where it can obtain the valley depth due to the keyhole effect. Whereas xCT may not measure the valley depth or any keyhole effect given that the xCT mesh may not be capable of exporting every deeply penetrated1763regions due to the keyhole parameter set. Sa shows a comparison where the highest difference is 34.17% for the EOS NOM parameter, respectively, where this parameter lies within the process window regime.

Sa shows a comparison where the highest difference is 34.17% for the EOS NOM parameter, respectively, where this parameter lies within the process window regime. For both Sku and Sz differences showed a comparison where the highest was 51.52% for Sku and 25.51% for the Keyhole parameter, respectively. Fig 7 shows each surface roughness value percentage difference between each surface roughness and printed parameter.5.ConclusionIn conclusion, the results that were displayed are partial and ongoing research is beingdone to further compare all technologies cohesively to further compare surface roughness as well as line profile roughness. While previous research has noted that surface roughness parameters such as average roughness (Sa or Ra) mean roughness depth (Sz or Rz), skewness (Ssk or Rsk) and kurtosis (Sku or Rku) could be used for the characterization of additive manufactured parts [4].

Our study and choosing of surface roughness parameters was done for the evaluation of fatigue tested samples where Sv could give us more insight into fatigue life and better understanding in the correlation of surface roughness and fatigue life. However, the correlation between fatigue life and surface roughness was not taken into consideration in this study.Current Sa, Sv, Sz, and Sku results do show a good comparison considering each printingparameter where Sa showed the highest difference of 34.17%, Sv showed the highest of differences up to 104.26 %, Sz highest difference of 51.52%, and Sku highest difference of 25.51%. Amongst the measured values, Sv showed the least comparable given the highest difference of 104.26% and the lowest difference of 18.71%. The reason for this is still unknown and with further comparisons, and a more thorough literature review, we strive to reach a conclusion and understanding.

The reason for this is still unknown and with further comparisons, and a more thorough literature review, we strive to reach a conclusion and understanding. Additional investigation will be done to have more insight and understanding why the keyhole regime had the highest values of Sv, Sz, and Sku. Factors that could affect these parameters can include but are not limited to scanning strategy, number of contours, as well as placement on the plate, where these factors were not taken into consideration in this study. Moreover, this study was not done to conclude what technology is best or more convenient but to give a better understanding of when to use a given technology for a certain and given application. For example, xCT will be of great use when the component being analyzed needs thorough internal and external investigation, given that xCT has the capability to internally and externally characterize a component that optical and profilometer methods cannot achieve.

Also standardizing xCT methods and having repeatability amongst surface roughness results will give the capability to implement this form of surface measuring method to complex geometries for industrial applications. Future research could be to compare these technologies and their surface roughness results for the qualification of manufactured parts. Doing so will give a representation of how these technologies could be used for the inspection and qualification of surface roughness for industry and research.17646.References[1] R. K. Leach, D. Bourell, S. Carmignato, A. Donmez, N. Senin, and W. Dewulf,“Geometrical metrology for metal additive manufacturing,” CIRP Ann., vol. 68, no. 2, pp. 677–700, 2019, doi: 10.1016/j.cirp.2019.05.004.[2] C. Sen, G. Dursun, A. Orhangul, and G. Akbulut, “Assessment of Additive ManufacturingSurfaces Using X-ray Computed Tomography,” Procedia CIRP, vol. 108, pp. 501–506, 2022, doi: 10.1016/j.procir.2022.03.078.[3] A.

Comparison of Synthesis Methods for Cellular Structures with Application to Additive ManufacturingJane Chu, Sarah Engelbrecht, Greg Graf, David W. RosenThe George W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332-0405 404-894-9668, david.rosen@me.gatech.eduAbstractCellular material structures, such as honeycombs and lattice structures, have been engineered at the mesoscale for high performance and multifunctional capabilities. We desire efficient algorithms for searching the large, complex design spaces associated with cellular structures. In this paper, we present a comparison of two synthesis methods, Particle Swarm Optimization (PSO) and least-squares minimization (LSM), for the design of components comprised of cellular structures. Computational characteristics of the algorithms are reported for design problems with hundreds of variables.

Computational characteristics of the algorithms are reported for design problems with hundreds of variables. Constraints from SLS and direct-metal manufacturing processes are incorporated to ensure that resulting designs are realizable. Two 2- dimensional examples are used to study the characteristics of the proposed synthesis methods.Keywords: cellular materials, lattice structures, design synthesis, particle swarm optimization, least squares minimization, additive manufacturing.1INTRODUCTION1.1 Cellular MaterialsThe concept of designed cellular materials is motivated by the desire to put material only where it is needed for a specific application. From a mechanical engineering viewpoint, a key advantage offered by cellular materials is high strength accompanied by a relatively low mass. These materials can provide good energy absorption characteristics and good thermal and acoustic insulation properties as well [8].

These materials can provide good energy absorption characteristics and good thermal and acoustic insulation properties as well [8]. Cellular materials include foams, honeycombs, lattices, and similar constructions. When the characteristic lengths of the cells are in the range of 0.1 to 10 mm, we refer to these materials as mesostructured materials. Mesostructured materials that are not produced using stochastic processes (e.g. foaming) are called designed cellular materials. In this paper, we focus on designed lattice materials.In the past 10 years, the area of lattice materials has received considerable attention due to their inherent advantages over foams in providing light, stiff, and strong materials [2]. Lattice structures tend to have geometry variations in three dimensions; some of our designs are shown in Figure 1. As pointed out in [6], the strength of foams scales as ρ1.5, whereas lattice structure strength scales as ρ, where ρ is the volumetric density of the material.

As pointed out in [6], the strength of foams scales as ρ1.5, whereas lattice structure strength scales as ρ, where ρ is the volumetric density of the material. As a result, lattices with a ρ = 0.1 are about 3 times stronger than a typical foam. The strength differences lie in the nature of material deformation: the foam is governed by cell wall bending, while lattice elements stretch and compress. The examples in Fig. 1 utilize the octet-truss (shown on the left), but many other lattice structures have been developed and studied (e.g., kagome, Kelvin foam). We have developed methods for designing lattice mesostructure for parts [17,20] and have developed design-for-manufacturing rules for their fabrication in SL.Methods of continuum mechanics have been applied to various mesostructured materials. Ashby and co-workers wrote a book on metal foam design and analysis [2]. They and others have applied similar methods to the analysis of lattice structures. The octet truss in Fig.

Ashby and co-workers wrote a book on metal foam design and analysis [2]. They and others have applied similar methods to the analysis of lattice structures. The octet truss in Fig. 1 has been extensively analyzed. Deshpande et al. [6] treated the octet truss unit cell as a collection of tension-compression bars that are pin-jointed at vertices and derived analytical models of their collapse behavior for many combinations of stresses. Their results match finite element model behavior well, but tend to under-predict the strength and stiffness of octet trusses due to their assumption of pin-jointed vertices. Wang and McDowell [21] extended this study to include several other lattice cells. Recently, we have been developing a more general analytical model of lattice behavior [11]. From our general model, models for octet and other lattice structures can be derived. We base our model on a single vertex with a collection of struts incident on that vertex.

From our general model, models for octet and other lattice structures can be derived. We base our model on a single vertex with a collection of struts incident on that vertex. This vertex model will be our base “unit cell” for representation and modeling purposes.1 1 10.5 0.5 0.50 0 0wd1 1 1d d db) skin with single layer of lattice structure b) skin with single layer of lattice structurew w w0.5 0.5 0.5uv0 0 0 0 0 00 0 0u u u0.5 0.5 0.51 1 1v v va) octet-truss a) octet-trussc) Skin with 2 layers of truss structure made in SL c) Skin with 2 layers of truss structure made in SL Figure 1. Octet-truss unit cell and example parts with octet-truss mesostructures.1.2 Design for Additive ManufacturingDesign for manufacturing (DFM) has typically meant that designers should tailor their designs to eliminate manufacturing difficulties and minimize costs.

However, the improvement of rapid prototyping, or Additive Manufacturing (AM), technologies provides an opportunity to re-think DFM to take advantage of the unique capabilities of these technologies. Several companies are now using AM technologies for production manufacturing. For example,Siemens, Phonak, Widex, and the other hearing aid manufacturers use selective laser sintering (SLS) and stereolithography (SL) machines to produce hearing aid shells, Align Technology uses stereolithography to fabricate molds for producing clear braces (“aligners”), and Boeing and its suppliers use SLS to produce ducts and similar parts for F-18 fighter jets. In the first three cases, AM machines enable one-off, custom manufacturing of 10’s to 100’s of thousands of parts. In the last case, AM technology enables low volume manufacturing and, at least as importantly, piece part reductions to greatly simplify product assembly.

In the last case, AM technology enables low volume manufacturing and, at least as importantly, piece part reductions to greatly simplify product assembly. More generally, the unique capabilities of AM technologies enable new opportunities for customization, improvements in product performance, multi-functionality, and lower overall manufacturing costs.

More generally, the unique capabilities of AM technologies enable new opportunities for customization, improvements in product performance, multi-functionality, and lower overall manufacturing costs. These unique capabilities include:Shape complexity: it is possible to build virtually any shape, lot sizes of one are practical, customized geometries are achieved readily, and shape optimization is enabled.Material complexity: material can be processed one point, or one layer, at a time, enabling the manufacture of parts with complex material compositions and designed property gradients.Hierarchical complexity: hierarchical multi-scale structures can be designed and fabricated from the microstructure through geometric mesostructure (sizes in the millimeter range) to the part-scale macrostructure.In this paper, we are interested in developing design methods that enable designers to take advantage of the shape complexity capabilities of AM processes.

Specifically, we develop design synthesis methods for cellular materials in general, and lattice structures in particular [17]. We first present a problem formulation for general structural synthesis applications (Section 2). In Section 3, we propose two synthesis algorithms, Particle Swarm Optimization and Least Squares Minimization, that show promise in searching the design spaces defined in our problem formulation. Two examples are investigated in Section 4 and the performance of the synthesis algorithms is compared.2 PROBLEM FORMULATIONThe class of problems being solved is most closely related to size optimization, although our problems have aspects of topology optimization and multiobjective decision making. Given the synthesis methods under investigation, we use unconstrained problem formulations, where constraints are included using penalty methods, if necessary.

Figure 3 is a math formulation of a problem to minimize volume and deviations from -/+ are under and overachievement deviation variables, target nodal deflections, where the di respectively, and Wi are goal weights (importances). For a given problem, we may or may not include the maximum stress constraint in the objective function.Various truss topologies can be explored by relaxing the lower bound on strut dimensions; when they approach 0 the strut is deleted. The notation x ∈ {0, [LB, UB]} indicates that strut dimensions can take on values between upper (UB) and lower (LB) bounds, or if they have values less than the LB, then they are set equal to 0.

Several design optimization methodologies could be applied to the problems of interest, including topology optimization, shape/size optimization, and hybrid methods. Topology optimization methods can be categorized based on their design representations and their search techniques. In the homogenization approach, a representation based on composite materials was used, where a material density function ρ models an infinite number of periodically distributed microstructures with small holes [3]. Therefore, the on-off material distribution problem is converted to a sizing problem and moves from the macroscopic scale to the microscopic scale. From a macroscopic perspective, a point in a structure can be partially occupied by the structural material, with ρ = 0 corresponding to a void, ρ = 1 to solid material, and 0 < ρ < 1 to the porous composite with voids at the micro level.

Many microstructures have been developed and generally fall into three categories: porous micro-microstructure [4], rank laminate composite [1], and free mixture [5]. However, this kind of on-off approach requires the use of discrete optimization algorithms, which can be unstable [9].In the discrete ground truss approach, the optimum topology is a subset of the ground truss, which is a complete graph of struts among all nodes. The cross-sections of the ground truss members are considered as continuous design variables for this optimization problem. The members with vanishing cross-sectional areas are removed to obtain the optimum [7]. Let a(i), l(i)denote the cross-sectional area and the length of beam number i in the ground truss shown in Fig. , and v(i) = 7. Its Young’s modulus is given as a(i) l(i) (i = 1,…,m) is beam volume for beam i as the fundamental design variables.

, and v(i) = 7. Its Young’s modulus is given as a(i) l(i) (i = 1,…,m) is beam volume for beam i as the fundamental design variables. For a typical single load situation, the design synthesis problem for the discrete ground truss approach is formulated as minimizing compliance, or deflection, and volume subject to static equilibrium and stress constraints [3,5].m( )iE . The given volume of the truss is= ∑i ( )Vvi1 =Recently, an exploratory framework was developed that can minimize the risk of structural failure by integrating a topology optimization method and a reliability assessment technique [15]. In this method, a Genetic Algorithm (GA) method [10] was used for the optimization process and Latin hypercube sampling was conducted for the estimation of the reliability constraint.

In this method, a Genetic Algorithm (GA) method [10] was used for the optimization process and Latin hypercube sampling was conducted for the estimation of the reliability constraint. Other researchers have applied GA methods for the synthesis of structural components, arguing that the evolutionary nature of GAs is well suited to the exploration necessary in the large complex design spaces typical of cellular materials [22].3.2 Particle Swarm OptimizationTo date, we have used a synthesis method based on Particle Swarm Optimization (PSO), which is an extension of genetic algorithms (GA), to perform parametric and limited topological optimization of structures and compliant mechanisms. PSO simulates the movement of birds in a flock, where individuals adjust their flying according to their experience and other individuals’ experiences during searches for food [12].

PSO simulates the movement of birds in a flock, where individuals adjust their flying according to their experience and other individuals’ experiences during searches for food [12]. It combines local search with global search, and enables cooperative behavior among individuals (“birds”), as well as the competition modeled using GA. Hence, PSO often converges more quickly than GA and was selected for the design synthesis of cellular structures [7].The search method of PSO creates a number of particles (the swarm), which “fly” in the design domain. Each particle updates its velocity and position according to its own experience as well as the swarm’s combined experience, according to Eqns.

The third term is a function of the best solution found by the entire swarm (pgd). The position update equation (Eqn. 5) is simply the sum of the current position of a particle and its velocity [12], where apparently it is acceptable to ignore the units mismatch in this community.Considerable research is needed to identify values of the four parameters that control PSO: the velocity inertia constant, wk, the cognition behavior constant ϕ1, the social behavior constant ϕ2, and the size of the swarm. In our experience, appropriate values are: wk = linearly interpolated from 0.9 to 0.4 through the PSO run, ϕ1 = ϕ2 = 2, and the swarm size is chosen in the range of 40 to 70 percent of the number of variables.3.3 Least-Squares MinimizationThe achievement of target values of goals can be formulated as a least-squares regression problem, which has similarities to formulations in inverse design [14] and parameter estimation(5)[18].

For cellular material design, the number of design variables far exceeds the number of objectives, which is similar to fitting a low order polynomial model to a large data set. The least- squares formulation for this problem is given by Eqn. 6.S() ∑X =(P i,target−P i actual ,()2 ) X(6)where Pi,target is the target value of the ith objective, Pi,actual is the actual value of the ith objective, and X is the vector of design variables. This error term is to be minimized, so the derivative of S is set equal to 0:∇S(X)=2n∑i1 =∂  P i actual , X ∂(X)    P i,target−P i actual ,(X) =0(7)where the partial derivative term is the Jacobian, J(X), of the system, Since J is nonlinear, an iterative solution technique must be used to solve for the unknown coordinates, X. Gauss- Newton methods are typically used to solve such problem.

Gauss- Newton methods are typically used to solve such problem. We used the Levenburg-Marquardt (LM) method [16], an extension of Gauss-Newton methods, since it tends to be more robust when sensitivities in the Jacobian are small. The iteration function for the LM method is:Xk+1 = Xk + [(Jk)T Jk + µkI]-1 (Jk)T [Pi,target – Pi,actual] where, µk is a scalar damping parameter that aids stability of the method.(8)We use Matlab to solve the process planning problems. Its non-linear least-squares solver,lsqnonlin, selects from Gauss-Newton and LM algorithms to solve problems.4 SYNTHESIS EXAMPLESTwo examples will be used to compare the performance of the PSO and LSM algorithms. One example is a cantilever beam composed of square unit cells, while the other is a simply supported bridge structure, which utilizes a ground truss as the starting configuration.

One example is a cantilever beam composed of square unit cells, while the other is a simply supported bridge structure, which utilizes a ground truss as the starting configuration. Both problems are in 2-dimensions.4.1 Cantilever Beam ExampleFour cantilever beam problems were investigated, each consisting of square 10x10 mm unit cells. The beams consist of 1x3, 3x8, 4x11, and 9x25 unit cells, where each unit cell consists of four beams (lattice struts) arranged in a square. As shown in Figure 4 for the 3x8 case, the left end is fixed and the right end is loaded with a 10 N point load. Design variables are the beam diameters. Target deflections of nodes at the free end are determined as 20 percent of the deflection of a solid beam (through finite-element analysis). Target volumes were: 226.2, 1407.4, 2448.1, and 11938 mm3 for the four cases.Figure 4 Cantilever beam problem, 3x8 case.Figure 5 Representative result of PSO for the 3x8 case.1x39 variablesTable 1.

Obj. Value2.3915 9340Objective Value 0.0099 0.0653Function calls 308 3109x25475 variablesPSORun #Init. Obj. Value -Objective Value -Function calls -Time [sec] -LSM Init. Obj. Value6.81 24000Objective Value 0.01 0.1412Function calls 958 1438Results are shown in Table 1. The table is organized by the sizes of the problems (e.g., 1x3, 3x8, etc.). Initial and final objective values are reported, along with the number of function calls to the objective function and the total time required. Multiple runs were performed for PSO since it is a stochastic algorithm. Note that the 15x40 case was too large to run and that only LSM could achieve a solution in the 9x25 case. The two LSM solutions in the 3x8, 4x11, and 9x25 cases represent problems where the initial designs had strut diameters of 2 mm and 1 mm,Time [sec] 0.609Time [sec] 15.92 26Time [sec] 102.9 103Time [sec] 7474 11745respectively.

Note that the 1 mm cases were initially far from optimum, improved significantly, but did not result in as low of an objective function value as the 2 mm case (except for 3x8).Results indicate that PSO and LSM achieve approximately the same objective function values, but LSM is one to two orders of magnitude faster than PSO. Example solutions for the 3x8 case are shown in Figure 5 and for the 9x25 case in Figure 6. Note that the PSO solution exhibits significant variations in strut sizes, but the variation does not follow obvious patterns. Although this may be expected due to the stochastic nature of PSO, we expected a more uniform decrease in strut size from left to right as PSO neared convergence.

A much more uniform size variation is observed in the LSM results.30282930Load Step NO: 1 3132333435362520192021222324252627151010111213141516171850123456789a)01020304050607080b)Load Step NO: 190235 236237238 239240241 242243244 245246247 248249250251 252253254 255256257 25825926080209 210211212 213214215 216217218 219220221 222223224 225226227 228229230 231232233 23470183 184185186 187188189 190191192 193194195 196197198 199200 201202203 204205206 20720860157 158159 160161162 163164165 166167168 169170171 172173174 175176177 178179 18018118250131 132133 134135136 137138139 140141142 143144 145146147 148149150 151152153 154155 15640105 106107 108109110 111112113 114115 116117118 119120121 122123 124125126 127128129 13030798081828384858687888990919293949596979899100101102 10310420535455565758596061626364656667686970717273747576777810272829303132333435363738394041424344454647484950515201234567891011121314151617181920212223242526050100150200250Figure 6 Results from LSM method for the a) 3x8 case, and b) 9x25 case.4.2 Bridge ExampleThe second synthesis example is a three-point bending lattice structure problem, as shown in Figure 7.

The premise of this example was to reduce the weight of the structure as much as possible without sacrificing stiffness at the central point load of –1 N. The bridge was assumed to be 40 mm long and 20 mm tall. This problem falls into the category of the well known Michell truss problems.Rather than utilize a unit lattice approach, we decided to start with a ground truss. Two versions of the problem were investigated, one with a 5x5 ground truss and one with a 7x7ground truss. Both problems were symmetric about the center-line; the 5x5 node problem had 102 variables while the 7x7 node problem had 377. As usual, the design variables are the beam diameters. It should be noted that the minimum value for strut diameters was constrained at 0.001, instead of zero, to maintain the mathematical stability of the finite-element code utilized to analyze the model.

After the optimization was complete, struts whose diameters fell below a lower threshold Dt=0.5 were removed from the structure, yielding a reduced number of individually sized struts. The target deflection of the top center node is zero. The relative weights of the deflection and volume goals are equal to 1.Initial experiments showed that both PSO and LSM had limited ability to identify appropriate solutions from a complex ground structure. To overcome this limitation, while preserving the routines’ access to the entire design space, an initial “seed” was presented at the onset of each problem. This seed consisted of a potential lattice design with elements that were likely to be appropriate for the final solution sized larger than those that were not.

This seed consisted of a potential lattice design with elements that were likely to be appropriate for the final solution sized larger than those that were not. The seed was placed as one member of the initial population for particle swarm optimization, while it represented the initial configuration of the system for least squares minimization.Figure 7 Ground truss (5x5) for Michell beam example.Again, LSM outperformed PSO by one or two orders of magnitude in terms of computation times as shown in Table 2. For the 5x5 case, PSO did not converge; synthesis was terminated after 200 iterations. In contrast, the LSM algorithm converged relatively quickly, using only 213 objective function calls, compared to 10,050 calls required by PSO. PSO did achieve a lower objective function value, 0.065 vs. 0.079 for LSM. Example solutions are shown in Figure 8.

PSO did achieve a lower objective function value, 0.065 vs. 0.079 for LSM. Example solutions are shown in Figure 8. Note that the PSO solution represents a very efficient, interesting solution, an isosceles triangle, which is close to the optimal solution.For the 7x7 case, PSO returned a wide range of solutions and converged sometimes. The best solution had an objective function value of 0.059, which required 10,050 function evaluations (no convergence). For the fastest convergence case, the solution had an objective value of 0.159 and required 2600 objective function evaluations. In contrast, LSM required 773 function evaluations to achieve an objective function value of 0.088. Representative solutions for the 7x7 case are shown in Figure 9. Note the resemblance to the optimal Michell truss [13].Figure 10 shows the typical progression of objective function values vs. generations for the PSO algorithm.

Note the resemblance to the optimal Michell truss [13].Figure 10 shows the typical progression of objective function values vs. generations for the PSO algorithm. Note that it is difficult to predict if or when reductions in objective function values will occur with this algorithm. We terminated PSO after 200 generations, but could have selected a different number of maximum iterations, which may have resulted in lower objective function values.Table 2.

We terminated PSO after 200 generations, but could have selected a different number of maximum iterations, which may have resulted in lower objective function values.Table 2. Results of PSO and LSM on the bridge problem.5x5PSOrun # GenerationsObjective ValueFunction callsLSMObjective ValueFunction calls1 2 3 4200 200 200 2000.164 0.696 0.065 0.06510050 10050 10050 100500.0792137x7PSOrun # GenerationsObjective ValueFunction callsLSMObjective ValueFunction calls1 2 3 4 5 640 200 100 200 25 2000.116 0.091 0.059 0.099 0.156 0.0844100 20100 10100 20100 2600 201000.0877772Figure 8 Example PSO (left) and LSM (right) solutions for the 5x5 problem.Figure 9 Example PSO (left) and LSM (right) solutions for the 7x7 problem.Figure 10 Typical progression of objective function values vs. generations for PSO.4.3 DiscussionBoth PSO and LSM produced results that were, for the most part, acceptable and appropriate for the problem.

generations for PSO.4.3 DiscussionBoth PSO and LSM produced results that were, for the most part, acceptable and appropriate for the problem. Particle swarm optimization was more apt to produce hanging struts, which are only attached to the structure on one end. This was most likely a result of the stochastic nature of the method, which introduces a certain amount of randomness into the solution. The differences in final objective function value between the two solution methods were on the order of 17% for the 5-by-5 truss, and 5% for the 7-by-7 truss, which is not significant for this portion of the design space. However, least squares minimization required approximately 96% fewer evaluations of the problem.

However, least squares minimization required approximately 96% fewer evaluations of the problem. For such complex problems, in which the time required for a single evaluation of the finite-element truss problem might be measured on the order of minutes, this reduction represents a significant reduction in processing time.While inclusion of an initial seed in the bridge problem dramatically increased the performance of both processes, its use reduces the extent to which the remaining portion of the design space is searched and requires pre-existing knowledge of the approximate solution. Both of these effects negate a certain amount of the utility of these processes, since computational optimization is implemented as a direct result of the large and complex design space.The need for an initial seed for LSM was not terribly surprising, since such gradient-based methods are often sensitive to the presence of local minima in the design space.

PSO’s difficulty locating an appropriate solution was unexpected, however, since the primary argument for its use is the ability to broadly search complex problems. It is possible that this difficulty could be alleviated through different settings of the various parameters of the optimization process (population size, particle velocity, etc). However, we used values that were recommended in the literature and fine-tuned by our experience. The difficulty in identifying more appropriate parameters lies in the length of time required for optimization, which is so great that it prohibits an exhaustive study of complex problems. If the parameters guiding the PSO process must be set individually for each design problem in order to provide accurate results, the usefulness of theprocess would be dramatically reduced.

If the parameters guiding the PSO process must be set individually for each design problem in order to provide accurate results, the usefulness of theprocess would be dramatically reduced. In the end, it might be more efficient to perform iterations of LSM and rely on multiple initial configurations to avoid local minimums present in the problem.Computational complexity results are displayed in Figure 11 as the number of objective function evaluations vs. the number of design variables. For the PSO runs in Tables 1 and 2, blue diamonds are displayed if PSO did not converge. For the two cases where PSO did converge (Table 2), pink squares are plotted. Green triangles indicate the results for LSM runs in Tables 1 and 2. It is clear that PSO takes many more function evaluations to achieve objective values that are comparable with those found by LSM.

Green triangles indicate the results for LSM runs in Tables 1 and 2. It is clear that PSO takes many more function evaluations to achieve objective values that are comparable with those found by LSM. The range of function evaluations appears to increase as the number of design variables increases, indicating a wide variation in execution times. When PSO does converge, considerable time savings are realized, but LSM converges more quickly. Although it is not clear from Figure 11, the LSM data points fall roughly on a linear curve, suggesting a very favorable linear increase in computational demands as problems get larger.2500020000PSO-No-Convs l a v E n F e v i t c e j b O #1500010000PSO-ConvLSM500000100200300400500# Design VariablesFigure 11. Number of objective function evaluations vs.

Number of objective function evaluations vs. number of design variables for the PSO and LSM methods.5 CONCLUSIONSTwo synthesis algorithms were compared for their use in designing structures with cellular materials, which are characterized by complex geometries and large numbers of design variables. Synthesis algorithms for such structures must be able to efficiently and effectively search large design spaces for promising design regions and for local optima. Two simple 2-D problems were explored, a cantilever beam and a simply-supported plate. When lattice materials are used to comprise the beam and plate, the design problems consist of hundreds of design variables.The objective function to be minimized was a weighted sum of deflection and volume.

Based on experiments with these problems and the Particle Swarm Optimization (PSO) and Least Squares Minimization (LSM) algorithms, the following conclusions can be offered:Both PSO and LSM found very good designs.LSM converged much more quickly than PSO, often by more than an order of magnitude fewer objective function evaluations.PSO was somewhat more effective in searching the large design spaces, as evidenced by the slightly lower objective function values that were found. However, PSO rarely converged in the range of generations that PSO was allowed to run.Good initial designs were needed to ensure good performance of both PSO and LSM.

However, PSO rarely converged in the range of generations that PSO was allowed to run.Good initial designs were needed to ensure good performance of both PSO and LSM. LSM needed a good initial design, while PSO needed at least one good design in its initial generation.LSM will be utilized for design synthesis in the future for sizing problems, while PSO will only be used when exploring large, complex design spaces where good configurations are not yet known.6 ACKNOWLEDGEMENTSWe gratefully acknowledge the U.S. National Science Foundation through grant DMI- 0522382, and the support of the Georgia Tech Rapid Prototyping and Manufacturing Institute member companies, particularly Pratt & Whitney and Paramount Industries, for sponsoring this work.7 REFERENCES [1] Allaire, G. and Kohn, R.V., “Optimal design for minimum weight and compliance in plane stressusing extremal microstructures,” Europ. J. Mech. A/Solids, 12, 6, pp.839-878, 1993.[2] Ashby, M.F.; Evans, A.; Fleck, N.

Comparison of the Computational Speed of 3DSIM versus ANSYS Finite Element Analyses for Simulation of Thermal History in Metal Laser SinteringKai Zenga,b, Chong Tenga,b, Sally Xub, Tim Subletteb, Nachiket Patilb, Deepankar Palab, Brent Stuckera,baDepartment of Industrial Engineering, University of Louisville b3DSIM, LLCAbstractA new simulation infrastructure for predicting the effects of changes in processparameters on mechanical properties, residual stress/strain, crystal structure, and other micro &macro features of components made using metal-based AM techniques has been developed at theUniversity of Louisville (UofL) and is being commercialized by 3DSIM, LLC. Based upon itsMatLab and Fortran code, UofL personnel predicted their multi-scale, multi-physics finiteelement solvers should solve for thermal history and residual stress evolution many orders ofmagnitude faster than competing tools while achieving better solution accuracy.

In order to testthis contention, a series of computational experiments were designed to benchmark theperformance of the code being commercialized by 3DSIM against a well-respected simulationtool, ANSYS. The results of these initial studies indicate the 3DSIM architecture is significantlyfaster than ANSYS for simulating metal-based AM processes.IntroductionA new simulation infrastructure developed at the University of Louisville [1-4] hasresulted in the formation of a new start-up company, 3DSIM, LLC. 3DSIM is creating acommercial implementation of these algorithms in C++ which will be highly optimized in thefuture to run efficiently in Graphical Processing Unit (GPU) based high performance computingenvironments. This paper presents an initial test of the speed of the 3DSIM code, prior to finaloptimization and release.

This paper presents an initial test of the speed of the 3DSIM code, prior to finaloptimization and release. As such it represents a first indication of the computational benefitswhich should be achievable with this code once fully implemented and optimized.In this work, the metal laser sintering processes is simulated using a moving meshstrategy to capture dynamic thermal fields during traverse of a Gaussian laser energy sourceacross the top of a powder bed. In the 3DSIM approach, a preprocessing architecture enablesintelligent stiffness assembly to predict the thermal fields in space and time. By adopting asimilar meshing strategy, a moving mesh model has been developed in parallel using thecommercially available software tool ANSYS [5, 6] with user defined mesh and boundaryconditions. Both models consider various heat transfer phenomena such as heat conduction,convection and phase changes that occur in metal laser sintering processes.

Both models consider various heat transfer phenomena such as heat conduction,convection and phase changes that occur in metal laser sintering processes. The 3DSIM modelalso considers residual stress/strain.ModelIn this study, the time to simulate the small part shown in figure 1(a) has been estimatedbased upon results for ANSYS and 3DSIM simulations of sub-portions of the part. The part wassimulated based upon standard process parameters for Ti6/4 used in an EOS M270 machine andthe use of a block support structure with a 2 mm height. The simulations were set up to calculatean addition 1 mm of unmelted powder on each side of the part and 1mm of the baseplate,resulting in an FEM model size of 42 mm x7 mm in the XY plane and an initial 3 mm Z height,which increases as new layers are added in the Z direction. The mesh used for this problem,showing a uniform coarse mesh throughout and a fine discretization near the energy source, isshown in figure 1(b).

The mesh used for this problem,showing a uniform coarse mesh throughout and a fine discretization near the energy source, isshown in figure 1(b). The laser beam diameter has been assumed to be 100 µm and is sub-divided into 8 elements. The coarse mesh discretization encapsulating the remaining geometryhas a 200 µm spatial grid size in the XY-plane. The layer thickness and the Z direction elementsize were assumed to be 30 µm. The Z-direction element size for the support structure and thebuild plate were set at 125 µm and the existence of the support structure was assumed to bealready present at time t=0 with properties representative of a block support structure. The initialstrip angle for scanning was observed to be 106.15° w.r.t. +X axis for an EOS M 270 machine,and this assumption has been replicated in figure 1(c).

The initialstrip angle for scanning was observed to be 106.15° w.r.t. +X axis for an EOS M 270 machine,and this assumption has been replicated in figure 1(c). All coordinates follow ASTM standardterminology [7].(a) A plate with dimensions of 40x5x2mm, built on top of a block support structure (in red) of2mm in height(b) Mesh for the structure(c) Scan strategy for the first layerFigure 1. Problem definition showing the part, mesh and scan strategy employed in the simulationResults and discussionA 3 dimensional model built in ANSYS using the above-mentioned discretization andoverall dimensions ranges from 204k to 714k degrees of freedom (DOFs). The number of DOFslinearly increase as the number of layers increase. Figure 2 shows the time to simulate varioussub-portions of the model for different layers on a Dell Precision T1650 desktop computer with aprocessor speed of 3.2GHz and 8GB of RAM.

Figure 2 shows the time to simulate varioussub-portions of the model for different layers on a Dell Precision T1650 desktop computer with aprocessor speed of 3.2GHz and 8GB of RAM. Each model ran for the first ten laser offsets foreach layer of simulation (it was not possible to finish running even one layer in an acceptableamount of time).(a) Running time versus number of substeps(b) Average time per substep versus DOFs (x105)Figure 2 ANSYS sub-model run time on an ordinary desktop computerThe way laser offsets are modeled are different in ANSYS versus 3DSIM. The dynamicmesh developed for this problem has a fine mesh region that is stationary with respect to thecoarse mesh for several time substeps as the laser energy traverses the fine mesh. Once the laserhas crossed a pre-determined location in the fine mesh, the fine mesh moves in the direction oflaser motion so that the melt pool remains within the fine mesh region.

Once the laserhas crossed a pre-determined location in the fine mesh, the fine mesh moves in the direction oflaser motion so that the melt pool remains within the fine mesh region. In the ANSYS model,this is approximated by causing the point of laser exposure to remain virtually stationary at thecenter of the fine mesh between time substeps until the laser position should have reached thefine mesh boundary (Figure 1b) leading to an advancement of the fine mesh by one coarse meshdistance in the direction of laser travel. In the 3DSIM model, the laser exposure changeslocation within the fine mesh region every time substep. The result of this difference is that3DSIM more accurately captures temperature variations. Since this experiment was designed totest the speed difference of 3DSIM versus ANSYS, no detailed analysis of accuracy was done.In figure 2a, the ANSYS execution time is plotted against the number of substeps fordifferent DOF scenarios.

In figure 2b, the average ANSYS time per substep is plotted against thenumber of DOFs. In order to predict the time needed for larger problems, the execution speed ofthe model using different computational configurations was tested and is shown in figure 3 fortwo cases of approximately 204k and 510k DOFs respectively. It was observed from these testresults that the model runs fastest using the combination of 1 GPU and 2 CPU cores. Thecorresponding results for average time spent per substep w.r.t. DOFs is shown in figure 4.In order to estimate the time for the execution of entire part model in ANSYS, the resultsshown in figure 4 have been for extrapolation. From figure 4, the time to solve increasingnumber of DOFs was assumed to be a linear function for DOFs exceeding 220k, and the datawas fitted using a linear function.

From figure 4, the time to solve increasingnumber of DOFs was assumed to be a linear function for DOFs exceeding 220k, and the datawas fitted using a linear function. Using this linear estimation methodology, the execution timefor ANSYS to fully simulating the small test peg model using 321,160 substeps (representing asolution of the laser for every 12.5µm of laser travel) would be about 61.5 years.Figure 3 Different computational configurations comparisonFigure 4 ANSYS model running time for multiple GPU processorsIn order to compare the computational speed of the ANSYS model with its 3DSIMcounterpart, a model with three elements in the Z direction was executed using 3DSIM. Theexecution time for finishing one layer with about 200K time steps are shown in table 1 for both3DSIM and ANSYS. A previous case study reported in SFF 2013 showed that the Matlab versionof 3DSIM’s algorithms was about 66 times faster than ANSYS [6].

A previous case study reported in SFF 2013 showed that the Matlab versionof 3DSIM’s algorithms was about 66 times faster than ANSYS [6]. From the results in Table 1,the initial (non-optimized) C++ 3DSIM model used for this study is about 294 times faster.Time steps Offsetss/time steps/offsetTotal(hours)3DSIM ANSYS200244 20058411779 N/A0.2 67.540.5 012.8 3763.2Table 1 – 3DSIM Processing Time for the model with 3 elements in Z directionConclusion and future workANSYS is arguably the best existing commercial tool for solving SLM thermal problems.Using ANSYS, however, to simulate even a relatively small SLM thermal problem requiresmany years of computational time.

The set of algorithms and software tools being developed by3DSIM show an initial advantage for speed, and thus capability, for simulating AM problems.Given meshing and other differences in algorithms between ANSYS and 3DSIM, results aredifficult to compare; but this discussion illustrates how a solver built for AM simulations usinghighly optimized algorithms and novel numerical solutions shows promise for significantlyreducing computational time.The 3DSIM solver used for this study, was an early version of the software. Subsequentto this study several enhancements have been implemented which significantly increase thespeed of the tool. In addition, unique algorithms developed by the co-authors and demonstratedon other problems, including intelligent Cholesky, Eigensolvers, DOF reduction, and periodicand higher order boundary condition solvers [8], have not yet been implemented into the 3DSIMC++ code.

in Proceedings of the Solid Freeform Fabrication Symposium. 2013. Standard, A., F2792. 2012. Standard Terminology for Additive Manufacturing Technologies. ASTM F2792-10e1. Deepankar Pal, N.P., Kai Zeng, Brent Stucker, An Integrated Approach to Additive Manufacturing Simulations using Physics based Coupled Multi-scale Process Modeling. Journal of Manufacturing Science and Engineering, 2014. In press..

mparison of Thermal Properties of Laser Deposition and Traditional Welding Process via Thermal Diffusivity MeasurementYu Yang, Omoghene Osaze Obahor, Yaxin Bao, Todd E Sparks, Jianzhong Ruan, Jacquelyn Kay Stroble, Robert Landers, Joseph Newkirk, Frank Liou University of Missouri RollaReviewed, accepted September 14, 2006AbstractLaser deposition is an effective process for mold and die repair. In order to improve the partrepair quality(cid:712)the process impact on thermal diffusivity and thermal conductivity needs to be understood for laser deposited, welded and virgin H13. In this paper, H13 tool steel samples were made by laser deposition, welding and virgin H13 and then cut into pieces. Experiments were conducted to investigate the thermal diffusivity and conductivity. A laser flash method is used to test these samples.

Experiments were conducted to investigate the thermal diffusivity and conductivity. A laser flash method is used to test these samples. The future work and opportunities are also summarized.Keywords: Laser Flash, laser deposition, tool steel, H13IntroductionNowadays, part repair technology is gaining more interest from military and industries due to the benefit of cost reducing as well as time and energy saving. Traditionally, part repair is done in the repair department using welding process. The limitations of the traditional welding process are becoming more and more noticeable when accuracy and reliability are required 1. In addition, the life of the mold and die parts after being repaired by welding process is much shorter than that of the virgin metal according to industrial experience. In LAMP (Laser Aid Manufacturing Process) lab, the laser deposition process is developed to repair the parts (Figure 1), for example the worn die shown.Figure 1.

In LAMP (Laser Aid Manufacturing Process) lab, the laser deposition process is developed to repair the parts (Figure 1), for example the worn die shown.Figure 1. The die after repair by laser deposition and machined by CNCThermal conductivity and thermal diffusivity play an important role in the life of molds and dies. Higher thermal diffusivity means that thermal equilibrium will be reached faster when the temperature changes. A good thermal diffuser will react more quickly to environmental446temperature changes. Higher thermal conductivity equates to the transfer of more thermal energy per unit of time under steady state conditions 2. In this paper, different samples of tool steel made by virgin, welding and laser deposition were prepared and the thermal properties of these samples were investigated by the laser flash method.1. Theory of Experiment.Laser flash method was first introduced by Parker, Jenkins etc in 19601.

Theory of Experiment.Laser flash method was first introduced by Parker, Jenkins etc in 19601. The front surface of asample is heated instantaneously and heat conducts through the sample. The back surface temperature T vs. time t has this relationship:tLT ),(=Q ρ CL⎡ + ⎢ 21 ⎣∞∑n=1− )1(n⎛ − ⎜⎜ exp ⎝π 22n 2 Lα t⎞ ⎟⎟ ⎠⎤ ⎥ ⎦(1) 3where Q is the radiant energy incident on the front surface at t=0, ρis density, C is specific heatand α is thermal diffusivity, L is the thickness of the sample. Two dimensionless parameters, V and ω can be defined:V (L, t)=T(L, t)/TM (2) 3/ Ltαπω=22(3) 3TM represents the maximum temperature of back surface. The combination of 1, 2 and 5 yields:+= 21V∞∑(-1)nexp(−ωn )2(4) 3= 1nEquation (4) is plotted in Fig. 2:Fig. 2. Dimensionless plot of back surface temperature history 3One way of determining α has been deduced from Eq (4) and Fig.2.

2:Fig. 2. Dimensionless plot of back surface temperature history 3One way of determining α has been deduced from Eq (4) and Fig.2. When V is equal to 0.5, ω is equal to 1.38, so447=α1388 .0 t 2/12 L(5) 3Where t1/2 is the time required for the back surface to reach half of the maximum temperature rise. Thermal conductivity λ is calculated using the value of thermal diffusivity and specific heat. λ= α⋅Cp⋅ρ (6) 3 ( α : Thermal diffusivity, ρ : Density, Cp: Specific heat, is determined by this laser pulse and almost equal to the energy of the laser)2. Experiment setup:The purpose of the experiment is to test t1/2 - the time required for the back surface to rise to half of the maximum temperature it can reach according to equation (5). The front face of a small disk-shaped sample was subjected to a very short burst of radiant energy. The source of the radiant energy is a laser irradiation.

The front face of a small disk-shaped sample was subjected to a very short burst of radiant energy. The source of the radiant energy is a laser irradiation. The resulting temperature rise of the rear surface of the sample is measured and recorded by the real time system.Figure.3 shows the experiment design. The experiment system contains laser, real-time system and temperature sensor. The laser pulse and data record are both controlled by real-time system. The laser was a 50 watt Diode laser with very short pulse at about 5ms. The temperature sensor is RTD (Resistance Temperature Detector) (Figure 4), in order to achieve good heat conduction from sample to the sensor, thermal grease was applied in between the sensor and samples. A thermal couple was used before as the temperature sensor. However RTD is more accurate and has been developed to be so small that it fit for very tiny sample.

A thermal couple was used before as the temperature sensor. However RTD is more accurate and has been developed to be so small that it fit for very tiny sample. The infrared (IR) sensor is also applied in many systems for recording the temperature rise curve 4. It can measure thermal properties at very high temperature. We chose RTD over an IR sensor for the reason that it is simpler and effective.Figure 3. Laser Flash Method Experiment DesignThe hardware and software of the Real-time System come from National Instrument and the program was based on the Lab-view 8. The real-time system can achieve two functions: control448laser to create laser pulse in desired duration, measure and record the back face temperature of samples through the RTD sensor.3. Sample preparation:Laser deposition cladding, welding cladding and virgin tool steel samples, which are all tool steel H13, were cut into small pieces with the size of 5mmX5mmX2mm.

Sample preparation:Laser deposition cladding, welding cladding and virgin tool steel samples, which are all tool steel H13, were cut into small pieces with the size of 5mmX5mmX2mm. The thickness of the samples was important for the test.Figure 4- Specimens4. Experiment:The sample was placed on the RTD sensor and the laser nozzle was aligned. The real-time system controls the laser to flash on the front side of the sample and record the rise curve of the back side temperature of the sample simultaneously. A typical temperature rise curve is shown in Figure-5. The value of t1/2 was measured as shown in Figure-5 and the thermal diffusivity and thermal conductivity was calculated according to equation 5 and 6. For virgin, welding and laser deposition tool steel samples, we assume all samples had the same density and specific heat: Cp=0.45 J/g-c, ρ=7.8X103 kg/m3.449Figure 5- typical back-face temperature vs. time rise curve5.

time rise curve5. Result:Six specimens were prepared for the experiment, two for virgin metal, two for laser depositionand two for welding. Each sample had three replicates. The results are shown in table 1:Table 1- Thermal Conductivity Experimental ResultsVirgin partWeldingLaser depositionSample No.#1 #2 #3 #4 #5 #6Dimension (mm)5.0X5.0 5.0X5.0 5.0X5.0 5.0X5.0 5.0X5.0 5.0X5.0Thickness (mm)2.011 2.028 2.008 2.055 2.039 2.012t1/2(s)0.102 0.098 0.173 0.185 0.167 0.145Average Thermal Diffusivity (X10-6m2/s) 5.50 5.83 3.23 2.83 3.67 3.875.673.033.77Average Thermal conductivity (W/mK)19.7 20.9 11.6 11.4 13.1 13.920.311.513.5The virgin samples can be regarded as the reference samples, and the thermal conductivity of virgin samples is in the range it should be - around 20 W/mK.

Both laser deposition and welding specimens have lower average thermal conductivity than virgin H13, welding drops about 43% and laser deposition drops about 1/3, but laser deposition has higher average thermal conductivity than welding by 22%.6. AnalysisThere are many reasons why the different sample types have different thermal properties. Laser deposition has small heat affected zone than conventional welding, so that laser deposition has more homogenous microstructure, it is one of the reasons why welding samples have lower450thermal diffusivity and conductivity than laser deposition samples.7. Conclusion:In this paper, the laser flash method to measure thermal diffusivity was introduced. The experiment system based on the laser flash method was designed and conducted.

Conclusion:In this paper, the laser flash method to measure thermal diffusivity was introduced. The experiment system based on the laser flash method was designed and conducted. The results show that repaired H13 tool steel parts have lower thermal diffusivity and conductivity than virgin metal and the parts repaired by laser deposition have higher thermal diffusivity and conductivity than those repaired by traditional welding. It means that laser deposition process will be better for part repair considering the thermal diffusivity and thermal conductivity than traditional welding process.The next step is to study the microstructure and compare the difference between laserdeposition, welding and virgin tool steel H13. Also the samples are tested at the room temperature. Thermal properties of the samples at higher temperature should be investigated and a furnace will be added on this system and to keep the samples at proper constant temperature.8.

Thermal properties of the samples at higher temperature should be investigated and a furnace will be added on this system and to keep the samples at proper constant temperature.8. Acknowledgement:This research was supported by the National Science Foundation Grant Number DMI-9871185, the grant from the U.S. Air Force Research Laboratory contract # FA8650-04-C-5704, and UMR Intelligent Systems Center. Their support is greatly appreciated.451Reference:1. Kunnayut Eiamsa-ard, Hari Janardanan Nair, Lan Ren, Jianzhong Ruan, Todd Sparks, and Frank W. Liou, "Part Repair using a Hybrid Manufacturing System", Proceedings of the Sixteenth Annual Solid Freeform Fabrication Symposium, Austin, TX, August 1-3, 2005.2. Douglas Veitch, Brush Wellman, “Making the Most of High-Performance Mold Materials”, http://www.moldmakingtechnology.com/articles/120505.html3. Parker, W. J., et, “A Flash Method of Determining Thermal Diffusivity, Heat Capacity, and Thermal Conductivity”, J. Appl. Phys.

****************************************************************************************COMPARISON OF SELECTIVE LASER AND ELECTRON BEAM MELTED TITANIUM ALUMINIDES L. Loeber1, S. Biamino2, U. Ackelid3, S. Sabbadini4, P. Epicoco2, P. Fino2, J. Eckert11Leibniz Institute for Solid State and Materials Research Dresden 2Politecnico di Torino 3Arcam AB 4Avio SpAAbstractIn the following paper we present the investigation of microstructure and mechanical properties produced by selective laser melting (SLM) and electron beam melting (EBM). The chosen alloy is a Ti-(46- 48)Al-2Cr-2Nb alloy which has a great potential in replacing heavy weight Ni-base superalloys in turbine blades. Cylindrical specimens were produced and characterized by optical microscopy (OM), scanning electron microscopy (SEM) and chemical analysis to determine the microstructure and composition.

Cylindrical specimens were produced and characterized by optical microscopy (OM), scanning electron microscopy (SEM) and chemical analysis to determine the microstructure and composition. In addition compression tests at room and elevated temperatures (700-800 °C) were carried out to identify the mechanical properties of the alloy.Introduction******************************************************************************************Titanium Aluminides (TiAl) alloys exhibit a wide range of interesting properties which makes them a good candidate for exchanging heavy weigh Nickel-based superalloys in jet turbine engines. They poses an excellent high temperature behavior in terms of strength, oxidation resistance, creep resistance and high melting point, while their density is only half as the one of Nickel [1-4]. Their poor room temperature ductility and their high reactivity with other materials makes them difficult to process [5,6].

Their poor room temperature ductility and their high reactivity with other materials makes them difficult to process [5,6]. Cost and time intensive methods like isothermal forging or special types of casting have to be used to produce useable parts.The process technology of additive manufacturing is known for polymers like Polyethylene for almost 30 years now. Only recently with the possibilities of high power lasers the technology has been adapted to metallic and ceramic materials [13]. This development lead to the additive manufacturing technologies of selective laser melting (SLM) and electron beam melting (EBM). The high energy source allows it to melt and micro weld metallic or ceramic powder. The principle work process of all additive manufacturing techniques is quite similar. Starting with a thin powder layer an energy source is focused on specific local areas of the powder layer. The energy source can be a laser, electron beam or in case of polymers a simple light source.

The energy source can be a laser, electron beam or in case of polymers a simple light source. The energy input has to be high enough to micro weld the powder to a substrate plate and to micro weld the different powder layers together. This three step iterative process can be seen in Figure 1. The contact to the substrate plate is essential as this contact ensures the mechanical stability but mainly to guarantee good heat dissipation. In a second step the building platform is lowered by the layer thickness and finally a coating device applies a new layer of powder and the process starts over again. This three step work principle is iteratively repeated until the desired part is layer wise built.The main benefit of additive manufacturing is the high amount of possible complex geometries. Almost every thinkable complex structure can be built. Examples are back tapers, cooling channels or special lattice structures.

Almost every thinkable complex structure can be built. Examples are back tapers, cooling channels or special lattice structures. Even structures which other processing technologies can not build are possible to build with additive manufacturing. Another advantage is the complete use of the material. Powder which is not melted into theparts can be sieved and used again. In contrast to subtractive manufacturing techniques where usually the process starts with a dense block and during the process material is taken away to create the desired part thus creating a lot of waste material like turnings. Additive manufacturing uses almost only the needed material for the parts and no waste material is created. As the selective laser melting is one of the younger process technologies the methods has only been applied to a clear number of materials like tool steels, stainless steels, aluminum alloys, Inconel, Cobaltchrome or some titanium alloys.

In this work the adaption to the material class of titanium aluminides is presented. As there are already some publications about electron beam melting of titanium aluminides and the techniques are quite similar to another this work tries to compare both techniques on this special material.As there are no existing process parameters for TiAl for SLM, an average between Titanium and Aluminum with the respect to the different heat conductivities, melting points and heat capacities for an orientation point was chosen for starting parameters. The following formula of the energy impact was used to produce different specimens with different energy impacts:E iP v x h Where Ei is the energy impact in J/mm³, P is the laser power in W, v is the scanning speed of the laser in mm/s, x is the hatching of the laser tracks in mm and h is the layer thickness in µm. A high energy input mostly leads to a higher density of the parts.

A high energy input mostly leads to a higher density of the parts. But with higher energy input the amount of residual thermal stresses also rises and as the room temperature ductility of TiAl is low a critical value should not be exceeded.Some additional remarks have to be given to enlighten the main differences between SLM and EBM while processing TiAl. Both names already gives account to the main difference of both processes: while at SLM the energy source to melt and micro weld the powder is a laser the energy source of EBM is a focused electron beam. This main difference in the general set up of the machines lead to additional differences during the actual building. While an electron beam of high quality needs a good vacuum, usually in the range of 10-3 Pa, the whole building space has to be set under vacuum during the building. The laser is much less affected by the atmosphere so SLM only works with an inert gas, normally argon or nitrogen, in special cases also helium.

The laser is much less affected by the atmosphere so SLM only works with an inert gas, normally argon or nitrogen, in special cases also helium. Working with a vacuum has some major benefits like a clean environment, which is especially important while working with reactive elements like titanium. But it also helps to outgas impurities incorporated in the metal powder and gives a good thermal insulation. Another benefit is the possible high scanning rate of the electron beam to preheat the powder. The EBM machine has the capability of heating the powder up to 1100°C. This advantage is quite crucial as the room temperature ductility and the fracture toughness of TiAl is quite low.

The EBM machine has the capability of heating the powder up to 1100°C. This advantage is quite crucial as the room temperature ductility and the fracture toughness of TiAl is quite low. The material exhibits a brittle-ductile transition between 700-800°C [10,11].Figure 1 Three step iterative process of additive manufacturing Materials and MethodsSimple cylinders with the dimensions of 5 mm diameter and 20 mm in height were created with a CAD program and then built in batches of 10 with a SLM 250 hl (SLM Solutions). The used gas atomized powder can be seen in Figure 2 a. The machine is equipped with a 400 W Yb-YAG fiber laser. The machine is also equipped with two oxygen sensors to ensure a low oxygen content during the building process as titanium is highly reactive with oxygen and titanium aluminides tend to lose their mechanical performance when they reach critical oxygen content [7,8].

Before starting the process itself the building chamber is flooded with 99,999% pure argon until a value below 0,02 % oxygen is reached. During the building of the specimens the building space was constantly flooded with argon ensuring a low oxygen content. With this process an oxygen content of around 100 ppm could be achieved. The specimens were directly built on a substrate plate without any further support structure which is normally used. The substrate plate had the same composition as the powder material: Ti-48Al-2Cr-2Nb and was micro pent before the process to ensure a good adhesion of the powder. The energy inputs of the specimens were chosen to be between 10-21 J/mm³. The laser power lied between 100-175 W, the layer thickness was set constant to 100 µm as the gas atomized powder has a powder size distribution of 45-100 µm and the layer thickness should not undercut the maximum powder size.

The hatching was chosen to be between 110 µm and 150 µm.The experimental conditions for building the EBM parts can be found elsewhere [9].Both types of specimens of EBM and SLM have been characterized via optical microscopy and electron microscopy to determine the microstructure. Additional characterization of the chemical composition and the porosity has been made. Both types of specimens have been analyzed twice by the different institutes. The chemical composition was assessed by means of inductively coupled plasma (ICP), infrared adsorption and/or thermal conductivity (LECO instrument) in both institutes. Specimens for optical microscopy analysis were prepared by polishing and etching in Kroll’s reagent, in order to investigate microstructure and porosity content. The porosity has been quantified according ASTM E 2109-01, while the amount of lamellar phase according ASTM E 562-01.

The porosity has been quantified according ASTM E 2109-01, while the amount of lamellar phase according ASTM E 562-01. The density of the SLM specimens were conducted via three types of density measurement: by simply measuring and scaling the specimens (geometric method), by the Archimedes method and by an area analysis of the polished samples. Mechanical testing in terms of compression test has been conducted with an Instron 8549. Additional compression tests at elevated temperatures have been carried out on the EBM specimens. For this measurement the cylinders have been electric eroded to smaller cylinders with a diameter of 3 mm and a height of 5-6 mm.100 µma100 µmFigure 2. a SEM picture of the used powder for EBM and b SEM picture of the used powder for SLMbResults and discussion The chemical composition of the base material of the SLM process, both powder types and both SLM and EBM specimens can be found in Table 1.

It can be seen that the amount of impurities of the EBM powder is not changed during the building. It can be assumed that the vacuum helps to keep the amount of pick-up of these elements down. Both the amount of oxygen and the amount of nitrogen lie beneath the critical amount of around 1000 ppm for oxygen and 600 ppm for nitrogen. The composition of the SLM material could already be conducted with the base material. The amount of impurities in the SLM material and powder are in the same range as the EBM powder. The amount of nitrogen and oxygen are a little lower in the starting material. As the gas atomization is conducted in inert atmosphere and with a crucible free method the amount of impurities during this process is kept to minimum. Only the amount of oxygen was determined in the SLM powder, as the oxygen content has the biggest influence on the mechanical properties.

Only the amount of oxygen was determined in the SLM powder, as the oxygen content has the biggest influence on the mechanical properties. It can also clearly be seen, that in contrary to the EBM process were no additional oxygen impurity arises during the process, the SLM process with its inert argon atmosphere leads to higher oxygen content in the SLM specimen. This is quite understandable, as even the purest argon still contains much more oxygen than a good vacuum. On the other hand can the vacuum of the EBM process lead to a higher evaporation of lighter elements such as aluminum. But the variation of the aluminum content of the EBM powder and specimen is around 0.3 wt.% and is thus not considered to have an influence. It is curious that the variation of Aluminum in the SLM specimens is further reduced by 0.1 wt% compared to the one of the EBM.

It is curious that the variation of Aluminum in the SLM specimens is further reduced by 0.1 wt% compared to the one of the EBM. The only explanation for this occurrence is that the energy input of the laser is not yet calibrated well enough thus leading to evaporation effects of the Aluminum. The chemical analyses were conducted on different places of the specimens to determine the homogenous distribution of the elements. The variation in the composition varied below 0.2 wt.% and is also considered to have no effect. The amount of the alloying elements such as Nb and Cr does not change.Table 1.

The variation in the composition varied below 0.2 wt.% and is also considered to have no effect. The amount of the alloying elements such as Nb and Cr does not change.Table 1. Chemical composition of the powder as received and the specimensElemental chemical composition in wt.% Al 34.1Cr 2.4Nb 4.8EBM powder (as received) EBM specimen SLM material (as received) SLM powder (as received) SLM specimen 32.333.42.25.132.92.64.832.72.64.82.64.8Fe0.030.030.056‐‐O0.060.060.0340.050.08N0.0040.0060.001‐‐C 0.0050.0070.004‐‐TibalancebalancebalancebalancebalanceThe amount and type of porosity has the biggest influence on the mechanical properties. The EBM process parameters have been adapted to produce a free macrospore material. In the EBM material only one type of small pores still occur caused by entrapped argon gas during the gas atomization.

In the EBM material only one type of small pores still occur caused by entrapped argon gas during the gas atomization. The amount of residual porosity lies below 2.0 % and can further be reduced via hot isostatic pressing to below 1.0 %. The porosity of the SLM specimens looks a lot different. As these are the first specimens build with this technology the process parameters are not fully optimized leading to a much higher residual porosity. Still two types of defects can be determined. Regarding Figure 3.a it can be seen that there are simply some regions where the energy input was not high enough to melt all the powder leading to big porous areas with partially molten or umolten powder around it reflected by the dark areas, especially in the outer region of Figure 3.b. On the other hand a great amount of cracks can be found in the whole specimen.

On the other hand a great amount of cracks can be found in the whole specimen. It can be assumed that the contrary to the too low energy input leading to the bigger pores in some regions the energy input seems to be too high leading to thermal tensions which are not tolerated by the material and thus resulting in thermal induced cracks. Figure 4 shows the relative density in dependence of the used energy input of the different specimens. A clear tendency to higherdensity with a higher energy input is visible for the SLM specimens. An asymptotic behavior can be observed. For the EBM specimens no value for energy input can be given. The difference in the three means of measurement for the density can be explained by the following reasons. The area analysis delivers at lower density even lower values of density. This effect is caused by the necessary metallographic preparation of the specimens which can cause nicks.

The area analysis delivers at lower density even lower values of density. This effect is caused by the necessary metallographic preparation of the specimens which can cause nicks. At higher values this effect is minimized as fewer defects cause fewer nicks. The Archimedes method delivers higher values of density over the whole spectrum as the geometric method. This effect is probably caused by the closed porosity which causes wrong buoyancy during the scaling in water. Overall it can be stated that the density of SLM parts can achieve a maximum density of around 97% which is muss less than their counterparts produced by EBM. A further optimization of the SLM process parameters have to be conducted in the future to improve the density.20 µma1 mmFigure 3: a Detailed SEM BSE picture of the SLM produced cylinder and b SEM overview of a whole cylinder cut in halfFigure 4.

Relative and absolute density in dependence of the energy input of the produced SLM parts via geometric, Archmides and area method and relative and absolute density of the EBM specimens before hippingThe surface of both EBM and SLM produced specimens can be seen in Figure 5. Both technologies are powder based and the energy input is localized on a small spot. This leads to partially molten powder which is fused with the surface. This effect is also responsible for the normally high surface roughness of EBM or SLM produced parts [14]. A post process like shot penning or simple grinding and polishing can be used to improvebthe surface roughness. In the picture of the SLM produced specimens horizontal surface cracks can be observed. Compared with Figure 3.b it is notable that at the outer region of the specimens the cracks are horizontal while in the inner volume the cracks are vertical to the building direction.

Compared with Figure 3.b it is notable that at the outer region of the specimens the cracks are horizontal while in the inner volume the cracks are vertical to the building direction. The horizontal outer cracks are probably caused by a wrong parameter choose for the contour. While the inner perpendicular cracks are caused by the effect the work principle of the technology that a hot layer is fused on a cold layer causing thermal and mechanical tensions. Both crack types are thus caused by too high thermal tensions which released in these cracks. The EBM produced surface on the other hand underlines the result from the density measurement that the parts are almost completely dense. No defects like open porosity or cracks can be found. A sort of repeating relief can be observed.

No defects like open porosity or cracks can be found. A sort of repeating relief can be observed. This results from the bigger layer thickness in the EBM process.500 µma500 µmbFigure 5: SEM picutre of the surface of the specimens a) EBM, b) SLM (the arrow indicates the building direction)Figure 6 shows the micorstructure of both EBM and SLM produced specimens. The EBM specimen (Figure 6.a) clearly shows a fine bimodal microstructure alternated to some zone of bigger equiaxed grains generally under the secondary electron (SE) contrast. The zones of bigger (but still below 50 µm) tipically follow the layerwise growing of the sample. In Figure 6 c this zones a mainly parallel to layers. The grain size is small which can be explained through the small size of the focused electron beam which leads to a rapid cooling. The SLM specimen on the other hand shows no clear microstructure under the SE contrast. No clear statement of the as-produced SLM microstructure can be made.

The SLM specimen on the other hand shows no clear microstructure under the SE contrast. No clear statement of the as-produced SLM microstructure can be made. However some features of the microstructure can be observed. Figure 3.a shows the micorstructure under back scattered electron contrast (BSE). In the upper part of the figure an extreme fine dentritic structure can be seen. This is the original microstructure of the gas atomized powder which is also a rapid cooling process. This feature reflects the powder particles which are only molten on the surface while the volume keeps its original microstructure. Through the middle of the picture a line of fine small collmular grains can be seen. This structure reflects the connection of different layers which can be described as the melt track. In addition some darker areas can be seen in the picture. These are either caused by minor fluctuation of the element content on the mico level or caused by different oriantations of the grains.

In addtion to this Figure 6.d) shows again the horizontal cracks at the outer region of the specimen and some pores with unmolten powder in it which underlines again the theory that the building parameters are yet to be optimized.100 µma100 µmb500 µmc500 µmdFigure 6: SEM picture of the microstructure of one of the a) EBM produced and b SLM produced cylinders and more detailed picture of the same c EBM specimen and d SLM specimen (the arrow indicates the building direction)500 µma500 µmFigure 7: SEM picture of the micortructure of one of the a) EBM produced and b) SLM produced cylinders after heat treatmentbAs no clear distinction in the microstructure of the SLM specimens could be made and thus no real comparison of SLM and EBM produced microstructure could be made, both types of specimens were treated with a heat treatment.

In Figure 7 the microstructure of both EBM and SLM specimens after the heat treatment of 2h at 1400°C (above the alpha transition temperature) can be seen. It can be observed that both samples have changed into the same fine lamellar microstructure. The pores caused by unmolten powder can still be observed in the SLM microstructure.Seven of the as-produced SLM specimens and eleven of the as-produced EBM specimens were mechanically tested under quasi static compression. Compression test were chosen instead of tensile testing because of the simpler testing geometry. For compression simple cylinders are sufficient while using tensile test complex tensile bars have to be produced or eroded out of a cylinder. The SLM specimens were only tested at room temperature as their high porosity already lead to poor results at room temperature.

The SLM specimens were only tested at room temperature as their high porosity already lead to poor results at room temperature. The EBM specimens were tested at room temperature (red curves; 2 specimens), at 700°C (blue curves; three specimens), at 800°C (rose and turquoise curve; three specimens) and at 850°C (green curves, three specimens). The poor results of the SLM specimens are not surprising when regarding the density values of the specimens of around 90%. The Young’s Modulus could be estimated to an arithmetic value of 50 ± 13 GPa which is around a third of the normal value. It is not surprising that the Young’s modulus is much smaller than expected as compression testing delivers normally a much smaller Young’s modulus than tensile testing. Also the ultimate compressive stress lies with a value of 612 ± 56 MPa far beyond normal results of titanium aluminides [12].

Also the ultimate compressive stress lies with a value of 612 ± 56 MPa far beyond normal results of titanium aluminides [12]. The compressive strain of 1.98 ± 0.55% would be a good result for tensile testing but compression testing normally generates values of up to 20-30%. The overall performance of the SLM parts in the mechanical testing can very well be explained through the poor values of density. Further studies of the mechanical properties are not performed until the density can be enhanced to higher values.The results from the EBM specimens then again show some really difference to the SLM specimens. At room temperature they achieve an ultimate compressive stress value of 1800 MPa which is in good agreement with other publications. The compressive strain of the samples however exceed the normal values of 20%. Also at higher temperatures the values of the compressive strain stay this high.

The compressive strain of the samples however exceed the normal values of 20%. Also at higher temperatures the values of the compressive strain stay this high. The values for the ultimate compressive stress at higher temperatures (see Table 2) lie beneath the typical values from the literature. This effect results probably from the small grain size which makes the material at higher temperature more vulnerable to the effects of creep. The compressive strain values stay mostly at the same high level. It has to be taken in account, that the value of 15% for 850°C is the arithmetic mean of 33%, 8% and 6%. So even at high temperatures some specimens still perform at high levels while others fail much earlier. The Young’s modulus at room temperature delivers the same value as the one of the SLM specimens.

The Young’s modulus at room temperature delivers the same value as the one of the SLM specimens. With higher temperature the value drops to 32 GPa.Table 2: Results from the compression tests of the EBM specimens at room and at higher temperature Temperature Ultimate compressive stress (MPa) (°C) 1800 RT 810 700 540 800 418 850Young’s Modulus (GPa) 54 42 32 39Stress at offset yield (MPa) 544 460 419 339Compressive strain (%) 40 31 33 15Compressive strain (%)Compressive strain (%)Figure 8 Results of the compression test a at room temperature of the as produced SLM specimens and b of the as produced EBM specimens at room temperature and at 700, 800 and 850°CSummary and Conclusion In this work we presented the differences between SLM and EBM produced cylindrical specimens of Ti-(46- 48)Al-2Cr-2Nb. Although both processes are fairly similar to each other the results from the two techniques vary.

Although both processes are fairly similar to each other the results from the two techniques vary. With EBM working under a vacuum and with preheated powder the results in density exceeds the one of SLM. EBM produced specimens showed an as produced density of around 98 % density with only small spherical pores. With additional hipping of the specimens the density can be enhanced to a value of >99 %. The microstructure of the EBM specimens exhibits a bimodal microstructure. An altering pattern of bigger (still smaller than 50 µm) and smaller (below 10 µm) grains can be observed. The mechanical properties of the EBM specimens under compression are comparable to those produced by conventional techniques. The SLM produced microstructure is difficult to determine as pores and cracks do lead to a lost in contrast under SE contrast. Under BSE contrast different features such as unmolten areas or melt tracks can be observed. The maximum achievable density up till now lies around 97 %.

Under BSE contrast different features such as unmolten areas or melt tracks can be observed. The maximum achievable density up till now lies around 97 %. But as the residual porosity is caused by cracks in the structure the mechanical performance under compression is poor.So fare the EBM process with preheating of the powder and the vacuum surrounding during building seems to fit better for the additive production of TiAl. In the future a main focus has to be set on the optimization of the process parameters of the SLM process and the implementation of a heater to reduce thermal tension during cooling.References[1] S. Herter: Spanbildung und Randzonenbeeinflussung beim Drehen intermetallischer Titaniumaluminide. thesis (2010) Technische Universität Berlin [2] Loria EA. Quo vadis gamma titanium aluminide. Intermetallics 2001;9:997-1001. [3] Appel F, Oehring M, Wagner R. Novel design concepts for gamma-base titanium aluminide alloys. Intermetallics 2000;8:1283-312 [4] I.

mparisons between Thermal Modeling and Experiments in Laser-Densified Dental Powder BodiesKun Dai, Xiaoxuan Li, Leon L. Shaw Department of Metallurgy and Materials Engineering Institute of Materials Science University of Connecticut Storrs, CT 06269, USAAbstractA three-dimensional thermal finite element model including the effect of the powder-to- solid transition has been developed to investigate the transient temperature distribution during laser densification of dental powder bed for the layer-by-layer fabrication. The model encompasses the effects of the temperature- and porosity-dependent thermal conduction and radiation as well as the temperature-dependent natural convection. The simulation result is compared with the experiments which establish the temperature dependence of the dental porcelain microstructure and utilize this dependence to construct the temperature distribution profile.

It is found that the trend of the simulation result matches the experiments very well.Keywords: Dental restoration, Laser processing, Finite element modeling, Thermal analyses, Powder melting and solidification.I. IntroductionSolid freeform fabrication (SFF) is an automated manufacturing process that builds three- dimensional complex-shaped structures layer-by-layer directly from CAD data [1]. The recent advancement of SFF has led to a multi-material laser densification (MMLD) process for dental restoration [2-4]. Through this approach artificial teeth are expected to be fabricated from a computer model without part-specific tooling and human intervention, thereby offering the potential to reduce the labor cost and increase the restoration rate.

However, as the densification of the dental porcelain powder is accomplished via a laser beam, an understanding of the temperature distribution in the laser-assisted densification process is necessary in order to understand the phase transformation and microstructural evolution of the dental porcelain powder during laser densification. The understanding developed will provide the guideline to optimize the laser processing condition so that the microstructure and thus the mechanical properties of the laser-densified bodies are similar to those obtained via traditional furnace annealing process. The approach of finite element modeling (FEM) has been taken in this study to investigate the temperature distribution during laser densification. The FEM approach is necessary because the relative density of the workpiece changes continuously with time during233laser densification until it reaches near full density.

The FEM approach is necessary because the relative density of the workpiece changes continuously with time during233laser densification until it reaches near full density. As a result, the thermal conductivity of the workpiece also changes continuously. Thus, the transient temperature field of the workpiece is too complex to calculate using analytical methods, and numerical simulation becomes necessary.Many numerical modeling efforts have been carried out to investigate the temperature field in various SFF processes [5-22]. However, they are not suitable for the laser-assisted dental restoration process which requires the numerical models to encompass at least the effects of the powder-to-solid transition, laser power density, closed-loop temperature control, temperature- dependent thermal convection, and temperature- and porosity-dependent thermal conduction and radiation. Most of the existing SFF models (e.g., Refs. 5 – 22) are not so sophisticated.

Most of the existing SFF models (e.g., Refs. 5 – 22) are not so sophisticated. Therefore, in this study we have developed a model that includes (i) a coupled-filed analysis between temperature and porosity fields, (ii) the incoming laser beam power with Gaussian distribution, (iii) the optical pyrometer simulation in addition to the closed-loop temperature control, (iv) powder-to-solid transition, (v) temperature-dependent thermal convection, and (vi) temperature- and porosity-dependent thermal conduction and radiation. The results from the numerical simulation have been compared with experiments and found to match the experimental measurements quite well.II. Model DescriptionThe model developed is shown in Figure 1 and consists of a dental porcelain powder compact with a dimension of 10mm length, 6mm width and 2mm height before laser densification.

Model DescriptionThe model developed is shown in Figure 1 and consists of a dental porcelain powder compact with a dimension of 10mm length, 6mm width and 2mm height before laser densification. In simulation the laser beam is modeled as a heat flux, Q, with a Gaussian power distribution as an incoming heat source, which moves at a constant rate (240 mm/s) along the X- axis as shown in Figure 1. The heat losses from the six surfaces of the model are assumed to result from thermal natural convection and the radiation. The dental porcelain powder is composed of 63.40% SiO2, 16.70% Al2O3, 1.50% CaO, 0.80% MgO, 3.41% Na2O, and 14.19% K2O (wt%). The material thermal properties for the dense solid porcelain used in the model are summarized in Table 1.

The material thermal properties for the dense solid porcelain used in the model are summarized in Table 1. The dental porcelain is assumed to be at a molten state above 1073K (100K higher than the lower temperature of the forming temperature range of the porcelain) and no volume shrinkage is considered when powder porcelain converts to dense porcelain.The modeling is carried out using the ANSYS commercial finite element package. The thermal element (Solid70), which has eight nodes with a single degree of freedom (i.e., temperature) at each node and has a 3D thermal-conduction capability, is used to simulate the temperature field [23]. Each element near the surface has a size of 0.25 mm length, 0.25 mm width and 0.2 mm height. The model is first used to calculate the temperature distribution in the powder bed within a small time step resulting from the heating of a laser beam moving at a constant rate along the X-axis.

The model is first used to calculate the temperature distribution in the powder bed within a small time step resulting from the heating of a laser beam moving at a constant rate along the X-axis. The powder elements convert to dense molten elements if their temperatures are higher than 1073K according to the calculated temperature field, and the temperature distribution in the powder bed within the next small time step is calculated using the updated material properties. This simulation loop continues until the total amount of the small time steps, which is decided by the laser scanning rate, is reached, and then the laser beam moves stepwise by one element to carry out the next simulation loop. The substrate preheating is234achieved in the simulation by assigning an elevated temperature (4600C) as the nodal temperature of the bottom surface.Heat FluxScanning RateRadiation ConvectionYXZPowder BedFigure 1.

It is assumed that the porosity of the powder bed is temperature independent before the powder becomes liquid. This is a reasonable assumption because the reduction in porosity due to solid-state sintering is minimal under the present laser densification condition which brings the local temperature of the area irradiated by the laser beam to above the forming temperature of the porcelain in less than 6 seconds. For the region outside the irradiated area, the time for the region to expose to high temperatures is also relatively short (less than 100 seconds) because of the scanning rate used. Furthermore, the temperature at the region outside the irradiated area is also relatively low because of the low thermal conductivity of the powder compact. As such, the porosity of material has been simplified in two levels, that is, the initial porosity j0 before the powder converts to liquid, and zero porosity (fully dense) after the powder23518002.82 1474 9.43 æhas converted to liquid.

The thermal properties of the powder bed, therefore, are treated as a function of temperature and the initial porosity j0. The details of these functions and thermal boundary conditions can be found in a recent paper [21] and described briefly as follows.A. Thermal Conductivity of the Powder BedEffective thermal conductivity of the powder bed, k, is estimated by Equation (1) with theassumption that the particles are spheres and there is no flattening of contact surfaces [29].

Ê Á Á Á Á Á Ëˆ ˜ ˜ ˜ ˜ ˜ ¯ˆ ˜ ˜ 1 ˜ ˜ ˜ ¯Ê Á Á Á Á Á Ëˆ ˜ ˜ ¯Ê Á Á Ëj 0 kk kk k1 kk2 kk kj 0++j - 0=+rrs1(ln1)((1))11ffffff11kkss Where, kf and k s are the thermal conductivities of the ambient air and solid particle, respectively; kr is the thermal conductivity portion of the powder bed due to radiation among particles and equals to [29,30]kr4 s=3xTFr(2)where sis the Stefan-Boltzmann constant, xr is the average diameter of the powder particles, T is the temperature of powder particles, and F is a view factor which is approximately taken as 1/3 [30].B. Thermal Radiation of the Powder BedThe emissivity of the powder bed, e, is obtained bye=e HHA1( -+A He ) S(3)where, AH is the area fraction of the surface that is occupied by the radiation-emitting holes, eS and eH are the emissivities of solid particle and hole, respectively.

For a powder bed composed of randomly packed, single-sized spheres, AH and eH are given by [31]HA=.1j 0 j 2 02.0 908 j - 2 0908+1(4)ande H=È Í Îj 0 j 0j 01e S+ .32082(È + .31 Í Îj 01e S082()2)2˘ ˙ ˚˘ +˙ ˚1(5)C. Thermal Convection around the Powder BedThe heat transfer coefficient, hc, is temperature and size dependent [32]236h c=kN u Lf(6)where L is the characteristic length of the specimen, Nu is the Nusselt number, and kf is the thermal conductivity of the fluid as defined before. Nu is given by [33]Nu=Nu0+È Í ÎPG / r r /5.0(1( P r+300 16/9 ))9/16˘ ˙ ˚6/1(7)when 10-4 < Gr Pr < 4¥1014, 0.022 < Pr < 7640, and Equation (7) are Grashof and Prandtl numbers, respectively [32].0uN = 0.67 for a plate [33].

Gr and Pr inGr=g3 Lbr - 2 TT ( f f h 2 famb)(8)andP r=Ch f p kf(9)where g is the gravitational acceleration, rf is the density of the ambient air, bf is the thermal volumetric expansivity and bf = 1/ T f for idea gases, hf is the viscosity of the air, Cp is the specific heat of the air. The effect of the variation of air properties with temperature is evaluated at Tf = 0.5(T +Tamb), where T is the surface temperature of the powder bed, Tamb is the ambient temperature [32].D. Thermal Boundary Conditions for Powder, Liquid and SolidThe part being built is assumed to be in contact with air and the heat loss through air is approximated through the natural thermal convection and thermal radiation between the part and the ambient air.

Since the model includes three kinds of material status (i.e., powder, liquid and solid), the thermal boundary conditions are very complicated and vary with porosity, phase status and surface temperature which is the function of incident laser power and laser scanning rate. Under the assumption of little convection of liquid within the molten pool due to its small size (~ 2 mm), liquid and solid have been assumed to have the same thermal convection boundary as the powder bed, i.e., thermal convection around the surface of liquid pool and solid is determined by the temperatures of the ambient air and the liquid and solid under consideration [see eqs. (6-9)].

(6-9)]. Heat loss qr due to radiation of solid and liquid is described by [32] )4( T= se s4qT(10)rambwhere T is the surface temperature of solid and liquid, and es is the emissivity of the dense porcelain, Tamb is the ambient temperature, and s is the Stefan-Boltzmann constant.237The model of a moving Gaussian distribution laser beam, pyrometer temperature simulation, closed-loop temperature control of the laser power are described in detail in the paper [20] and is described as follows.E. Heat InputThe heat flux of the laser beam, Q, is related to the power of the laser beam, P, throughthe following relation [34]:Q =aP 2 p r 0a e222 r 2 r 0(11)wherea is the absorptivity of the workpiece,a0r is the radius of the laser beam at which the heat2e - times of that of the laser beam center, and r is the distance of a point on theflux value is surface of the powder bed measured from the laser beam center.

When the center of the laser beam scans the surface of the powder bed from the starting point (X0, Y0, Z0) to (X, Y0, Z0) point along the positive X-direction with a constant velocity V for a period time t, r is given byr2=(XX0Vt)2+(ZZ 0ZZ 02(12)In the present study, the radius of the incident laser beam,0r , is 0.5 mm, the lasera , was scanning velocity, V, is 0.24 mm/s. The absorptivity of the dental porcelain workpiece, varied in the simulation, and a value of 0.41 was needed for the simulated laser input power to match the experimental values. This absorptivity value is comparable with experimental scatter for measured values ofa for sintered SiC near the processing temperatures used here [35].aaF. Pyrometer SimulationIn laser-densification experiments, an optical pyrometer continually monitors the temperature distribution at the surface of the powder bed during laser densification.

Pyrometer SimulationIn laser-densification experiments, an optical pyrometer continually monitors the temperature distribution at the surface of the powder bed during laser densification. This pyrometer temperature is used as the feedback signal in a closed-loop control program to adjust the incident laser power as needed to achieve a desired constant laser spot temperature.

In the simulation, the pyrometer temperature measurement and the closed-loop control process are modeled as follows.The power, E, of the thermal radiation emitted by the laser-heated workpiece that reachesthe pyrometer can be expressed by [35]E= ÚIl D ,( ) TldA(13)where l is the wavelength of the emitted radiation , lD is the wavelength band of the emitted radiation that is sampled by the pyrometer, T is the temperature at a very small area dA through which the radiation passes to reach the pyrometer, termed the pyrometer sampling area hereafter, I l is the spectral distribution of blackbody emissive power and given by Planck’s radiation ,( T law [36]:)238Il ,(T)=l 5p 2 Ê Á Á Ëhc hc ls Te2ˆ ˜ 1 ˜ ¯(14)where h is Planck’s constant, c is the speed of light, and s is Stefan-Boltzmann constant. Because of the Gaussian heat input and movement of the laser beam, the surface temperature, T, I l.

Because of the Gaussian heat input and movement of the laser beam, the surface temperature, T, I l . Thus, to carry out the ,( T within the pyrometer sampling area is not uniform, and neither is integration of equation (13), approximations are made by dividing the sampling area into n small areas and assuming that T and I are constant within each small area. With these assumptions, equation (13) is reduced to)E=nÂIl ,(T i)l DDA i(15)i= 1where Ti is the surface temperature of the area DAi and n is the number of the small areas within the pyrometer sampling area, Ap.

To further simplify the computation, it is assumed that every DAi has the same area, and thusApD=An i(16)and equation (15) becomesEl DD=A inÂI,(lT i)(17)i= 1To relate the thermal radiation power collected by the pyrometer to the pyrometertemperature reading, an effective temperature, Teff, is introduced asElDD=InA il ,(Teff)(18)Combining equations (14), (17) and (18), it hasTeff=lsÊ Á 1ln Á Á Ë+n/hcnÂi= 1Ê Á Á Ëehc ls T i1ˆ ˜ ˜ ¯1ˆ ˜ ˜ ˜ ¯(19)G. Closed-Loop Temperature Controlin Equation (13) is the effective (nominal) surface temperature of the workpiece as measured by the pyrometer for a given incident laser power. If Teff differs from the desired laser spot temperature, Tp, then the incident laser power is adjusted accordingly to simulate the closed- loop temperature control in the experiment.

If Teff differs from the desired laser spot temperature, Tp, then the incident laser power is adjusted accordingly to simulate the closed- loop temperature control in the experiment. This is achieved in the simulation by adjusting the laser power from one simulation step to the next using the following equationTeff239P =+1iP iTpTeff(20)where Pi and Pi+1 are the incident laser power in the simulation steps i and i+1, respectively. It can be found that this equation allows the surface temperature of the workpiece, Teff, to quickly approach the desired laser spot temperature, reproducing in the simulation the closed-loop temperature control process used in the experiments.III. Comparison between Simulation and Experimental ResultsOne of the key criteria for validating the model is to check whether the model can achieve a constant pyrometer temperature (i.e., a constant temperature on the surface of the powder bed) by continuously adjusting the laser input power.

Figures 2 and 3 show the comparison between the experimental and simulated pyrometer temperature as a function of the location of the scanning laser beam with a nominal surface temperature of 900 and 10500C, respectively. It can be seen that the variations in temperature are less than 4% of the desired pyrometer temperatures for all four nominal surface temperatures modeled, which indicates that the closed-loop temperature control in experiments can be simulated with the present model. Furthermore, the pyrometer temperatures achieved by simulations and experiments match quite well.)C e e r g e D12001000(r e t e m o r y p f o e r u t a r e p m e T8006004002000SimulationExperiment012345Location of laser beam (mm)Figure 2.

Comparison between the experimental and simulated values of the pyrometer temperature as a function of the location of the scanning laser beam with a nominal surface temperature of 9000C and substrate preheating to 4600C.The model established is utilized to predict the temperature distribution in the porcelain body during laser densification. Figure 4 show the simulation result with a nominal surface temperature of 10500C and substrate preheating of 4600C. It is quite clear that the temperature distribution obtained is consistent with the expectation that the highest temperature is located at240the center of the scanning laser beam and the temperature gradually decreases as the location moves away from the center in all directions within the porcelain body. More importantly, when compared with the experimental result (Figure 5), it is found that the predicted temperature distribution pattern matches the measurement quite well.

More importantly, when compared with the experimental result (Figure 5), it is found that the predicted temperature distribution pattern matches the measurement quite well. Furthermore, the temperature range predicted also matches the measurement reasonably well. However, a direct comparison of the temperature value between the simulation and experiment is not possible at this stage because of the different laser densification conditions used in Figures 4 and 5. The work on simulations with the identical laser densification condition as the experiment is currently under way and will be published in a forthcoming paper [37].)C e e r g e D12001000(r e t e m o r y p f o e r u t a r e p m e T8006004002000SimulationExperiment012345Location of laser beam (mm)Figure 3.

Comparison between the experimental and simulated values of the pyrometer temperature as a function of the location of the scanning laser beam with a nominal surface temperature of 10500C and substrate preheating to 4600C.Figure 4. Temperature distribution in the powder bed at the cross section of x = 6.75mm when the laser beam scans to this location with a nominal surface temperature of 10500C and substrate preheating to 4600C.241~ 1100~ 13000C~ 15000C~1100~12001100~12001100~1200~ 8001100~1200~ 800>1200~ 800~ 800~ 800500umFigure 5. Temperature distribution at the cross section of a porcelain body perpendicular to the laser scan direction during laser densification with a nominal surface temperature of 11000C and substrate preheating of 4000C [37].IV.

Conclusions and RemarksA 3D finite element model that encompasses the effects of the powder-to-solid transition, laser power density, closed-loop temperature control, temperature- and porosity-dependent thermal conduction and radiation as well as the temperature-dependent natural thermal convection has been developed to carry out the thermal analysis of laser-densified dental porcelain bodies. The temperature distribution in the porcelain body during laser densification has been simulated using this model and is compared with the experiment. The results predicted by the model matches the experiment well, and the model will be used in the future to provide guidance for selecting laser processing conditions to obtain the desired microstructure and geometry of the dense body. This is the first comprehensive model that includes all key parameters in laser densification of powder lines.

This is the first comprehensive model that includes all key parameters in laser densification of powder lines. The model can be extended to simulate many other laser materials processing methods that require consideration of phase transformations, closed loop temperature control, and heat transfer.Acknowledgements – The authors gratefully acknowledge financial support provided bythe National Science Foundation under Grant Nos: DMI-9908249 and DMI-0218169.242References1. Beaman, J.J., Barlow, J.W., Bourell, D.L., Crawford, R.H., Marcus, H.L. and McAlea, K.P., Solid Freeform Fabrication: A New Direction in Manufacturing, Kluwer Academic Publishers, MA, 1997.2. L. Shaw, X.-X. Li, J.-W. Wang, H. L. Marcus, T. B. Cameron, and C. Kennedy, “Dental Restoration through Laser Densification of Dental Porcelain Powder,” in Rapid Prototyping of Materials (F.D.S. Marquis and D.L. Bourell, Eds.), TMS, Warrendale, PA, 2002, pp. 107- 118.3. J.-W. Wang, X.-X. Li, L. Shaw, H. L. Marcus, T.

MPENSATION FOR UNEVEN SURFACES WHEN BUILDING LASER DEPOSITED STRUCTURESE. Fearon, K.G. Watkins University of Liverpool Laser Group Department of Mechanical Engineering, Brownlow Street, Liverpool EnglandReviewed, accepted August 31, 2004AbstractDirect Laser Deposition (DLD) is a blown-powder laser deposition process that can be used to quickly produce, modify or repair fully-dense metallic parts by a layered manufacturing method. However, uneven substrate surfaces often cause variation in the deposited layer which is magnified by succeeding layers. Research carried out at the University of Liverpool has resulted in a non- feedback layer height controlling process based on controlling the shape of the powder streams emitted from a four-port side feed nozzle.

This method limits deposited layer height by causing a sharp reduction of catchment efficiency in the vertical plane at a fixed distance from the powder feed nozzle, and is therefore capable of depositing a consistent layer height in spite of power, powder flow or process velocity variation. This paper demonstrates how this method of layer height control can compensate for irregular substrate surfaces in the production of accurate DLD parts.1. IntroductionDirect Laser Deposition (DLD) is an extension of the laser cladding process in that it allows three- dimensional parts to be built by cladding successive layers on top of one another in pre-determined vector paths (work began on the process in Liverpool in 1990 with the first paper published in 1993 [1]). Since the layers are fusion bonded to each other, a fully dense metallic part can thus be made using ‘soft tooling’ (i.e.

Since the layers are fusion bonded to each other, a fully dense metallic part can thus be made using ‘soft tooling’ (i.e. in order to make a different part only the program used to control the CNC equipment needs to be changed rather than needing to change tools or make new moulds).The projected usage for DLD include reducing development cycle time by rapidly manufacturing functional prototypes, modifying or repairing high- value existing components, and making components with inbuilt voids for weight reduction [2]. Projected uses for the process include the manufacture of spare parts for long term space missions [3] or on board ships or submarines [4].DLD is a multivariable process with interdependence of laser power, powder feed rate into the melt pool, relative traverse speed of the workpiece and retained heat effects causing difficulty in retaining a consistent clad bead profile for the purposes of building or modifying parts.

Power control has been extensively researched but necessitates in-process monitoring and feedback control in addition to the necessity to calculate (and account for) projected alterations of power density and interaction time for specific geometries.471In addition, when a clad bead is laid onto a previously produced layer, a number of extra considerations in addition to the utilised process parameters affect its generation. One of the most important of these is the effect of elevated temperatures in this previous layer due to insufficient cooling time or heat retention in the bulk of the part. This effect was recognized by Weerasinghe (1985) [5] and has been examined in greater detail by Vasinonta, Beuth and Griffith (1999) [6] who produced an ABAQUS generated model using Rosenthal (1946) [7] conditions which shows that the melt pool length and height increases as the initial temperature of the substrate is increased.

Therefore, unless steps are taken to control the layer height during a build, the retained heat in the part from previous layers will result in a cumulative increase in the melt pool volume, and hence the height and width of the deposited layer. In addition, this variation of melt pool size affects cooling rates and hence microstructure of the finished part. Alterations in process velocity as the CNC table or flying optics assembly change direction will also cause a change in deposited layer height, which on successive layers will again be magnified [8]. For these reasons deposition is normally carried out on flat surfaces or the surface is carefully modelled before commencement to prevent unwanted alteration of the deposited layer heights.In the repair and/or modification of existing parts uneven surface profiles may necessitate rapid alteration of the height of the deposited layer in order to compensate.

This would normally require extensive modelling of the surface or extremely responsive feedback systems.Research carried out at the University of Liverpool has resulted in a method of controlling the deposited layer height by abruptly limiting the availability of powder in the vertical plane at a fixed point relative to the powder feed nozzle [9]. This self-limiting system enables laser deposition to be carried out on irregular surfaces, as will be illustrated in the following sections.2. The method of layer height control.The nozzle system designed to control the deposited layer height appears below in Fig. 1. It consists of four radially symmetric 2mm bore powder feed tubes at 40 degrees to the horizontal. A small-bore (2mm) gas nozzle coaxial with the laser beam is used to direct a gas flow at the area where powder streams emitted from the powder feed nozzles will intersect.Fig.

A small-bore (2mm) gas nozzle coaxial with the laser beam is used to direct a gas flow at the area where powder streams emitted from the powder feed nozzles will intersect.Fig. 1 The four port nozzle systemPowder is delivered from the powder feed tubes at a relatively low velocity (<1m/s) and flares as it exits. Due to stratification of differently-sized powder particles (in addition to the Magnus effect for irregular particles)[10], powder flaring from the top of the powder stream as it exits the powder feed tube consists of smaller particles than the bulk flow. This, coupled with the fact that particles travelling off-axis from the main stream necessarily have lower particle velocities, results in the fact that powder flaring from the top of the powder stream has a lower momentumthan that in the main part of the powder stream. (Fig. 2)472Fig 2.

(Fig. 2)472Fig 2 . A false colour image showing the powder stream exiting a 2mm bore powder feed tube at a 400 angle under a gas flow rate of 4 l/min without the effect of a coaxial gas flow. The different colours indicate relative density of powder stream based on degree of reflected light. It can be seen that areas of lower powder density at the top of the powder flow is greater than that at the base. A ‘raw’ photograph of the same feed tube is shown for comparison on the right.A gas flow coaxial to the laser beam can thus be used to redirect these off-axis powder particles back into the main powder stream. Experiment has shown that a ‘balance point’ between the powder assist gas flow rates and the coaxial gas flow rate can be obtained (dependent upon the density and size distribution of the powder) which will result in a refinement of the top interface of the powder streams[9]. An example of this effect is shown in Fig.

An example of this effect is shown in Fig. 3.This refinement of the powder cloud interface causes an abrupt reduction in powder catchment efficiency in the vertical plane at a fixed position relative to the powder feed nozzle. In other words, a deposited layer will build to the top of this interface but can not build any higher due to a lack of available powder. This means that the height of the deposited layer is limited by the physical position of the powder feed nozzle and not by the other deposition parameters, allowing the layer height to be controlled by the incremental step height rather than vice versa.

In addition the low particle velocity assists in the improvement of catchment efficiency[5], offsetting the powder loss that would be a side effect of programming a step height which is less than the unrestricted layer height.Assessment of the capabilities of this method of layer height control has shown that it is capable of depositing layer heights controllable to within 300 µm of the desired layer height and that height errors are not cumulative in the build. Excess laser power, increased powder feed rate and reduced process velocity do not cause the deposited layer height to increase in proportion [11].Since the basis for redirection of the top of the powder is due to momentum differences of particles within the powder stream it is necessary to obtain a new ‘balance point’ for different powder material densities.

An increase in powder particle density will require an increase in the gas flow rate required for adequate transport of the powder through the powder feed tubes, in addition to an increase in the coaxial gas pressure required to redirect off-axis particles. At the473time of writing, more analysis using different materials needs to be carried out to generate the appropriate relationship .Fig 3. These two images show the effect of the use of a 12 l/min gas flow through a 2mm coaxial gas nozzle situated 15mm above the intersection of the powder streams. It can be seen that the powder stream on the right has had the top ‘penumbra’ reduced increased vertical and shows an collimation below the intersection point of the streams. (Measurement indicated by dashed line equivalent to restriction point of a 2mm wide layer)3. Control of layer height on regular surfacesFig.4 Square structures built with different layer heights under constant deposition parameters.

Control of layer height on regular surfacesFig.4 Square structures built with different layer heights under constant deposition parameters. Eachsquare is 2cm on a side.A simple square structure was used to evaluate layer height control. The abrupt change in velocity as the CNC table alters direction at the corners would normally result in a cumulative height error at the corners. To test the ability of the system in layer height restriction further, a set474of square structures (Fig.4) was built with identical process parameters excepting that the step height between layers was progressively reduced between samples. Under these conditions it was shown to be possible to control the deposited layer height to an error of approximately 300 µm. This 300 µm error in height was not cumulative, i.e. increasing the amount of layers present in the part did not increase the overall part height error. A graph of these results is shown (Fig.

This 300 µm error in height was not cumulative, i.e. increasing the amount of layers present in the part did not increase the overall part height error. A graph of these results is shown (Fig. 5).Ideal height of partHeight of part504540)m m35(t r a p f o t h g i e H30252015105000.511.52Input step height (mm)Fig.5 Actual height of deposited structures against desired height over a range of layer heights. (The falloff in part height beyond 1.3mm step height is due to the programmed step height being more than the possible clad layer height.)4. Control of layer height on surfaces with raised or recessed featuresOne of the major uses for laser deposition is in the modification of existing parts which may need an extra feature added for purposes of strength, wear resistance, extra attachment points or repair/reinforcement.

The surface to be modified may have raised or recessed features which would normally require surface profiling carried out in order to allow program modifications to be made. A specimen example of parts which would fall into this category appears in the series below (Fig. 6a-d). In Fig. 6a an irregular surface is shown that would normally require modelling to adjust the deposition parameters accordingly within the program, or a highly responsive feedback system. This surface has step, semicircular and angled cutouts to a depth of 3.5mm. Utilizing the layer height control process detailed in the previous section the deposition process for a simple wall is carried out such that the first layer is built corresponding to the lowest point on the irregular surface (Fig. 6b). Deposition is not carried out on surfaces higher than this until the gradual raising of the nozzle brings them into the defined area of the powder cloud. Fig. 6d shows the repaired surface which has been machined.

Fig. 6d shows the repaired surface which has been machined. No cracks or gaps are evident.475Fig. 6aFig. 6bFig. 6cFig. 6dFig. 6 a-d. Stages in repair of an uneven substrate surface.5. Incorporation of new components.One of the interesting features of being able to deposit structures upon irregular surfaces in this manner is the ability to incorporate small components into the part being built or modified. Fig.7 shows a structure built at Liverpool where a stock mild steel strip has been incorporated into a deposited billet. This strip has a thickness comparable with the laser beam spot size and as a result has been laser formed (i.e. bent using the temperature gradient mechanism) at the same time. The degree of layer height control has been sufficient to account for this forming process to be accommodated within the build process.

The degree of layer height control has been sufficient to account for this forming process to be accommodated within the build process. A further example of this incorporation process currently being investigated at Liverpool is the incorporation of cooling channels into a deposited or pre-existing part by simply building over pre-shaped metal piping rather than programming voids in the structure. Advantages to this are that the cooling channels incorporated in this manner are smooth on the inside without any generated layer discontinuities in the channel wall, and thus can utilize higher fluid flow velocities without generation of cavitation within the coolant fluid. A further use proposed is in the generation of quasi- hollow structures by building parts around pre-manufactured hollow pre- forms, thereby avoiding the conditions conducive to crack propagation under cyclic stresses noted by Capshaw [2] .Fig 7. Laser formed part incorporated into deposited billet.4766.

Laser formed part incorporated into deposited billet.4766. Retained heat and microstructure effectsControlling the deposited layer height in the manner described causes deeper remelting of previous layers than would be the case for unrestricted layer deposition or power feedback layer height control methods. This results in lower cooling rates and increased coarsening of the as- deposited microstructure.

Fig.8 illustrates this effect. This part was built by building a cylinder on a rotating substrate such that the outer layers were taller than the inner ones. This allowed a build where reheating frequency, velocity and power were constant.67mm27mmFig 8. Axially curved DLD cylinder (sectioned)Examination of Fig.8 shows that the inner (27mm radius) clad wall with a layer height of 0.32mm increases in thickness to a much greater degree than the outer one with layer height of 0.8mm. From an initial wall width of 2.2mm on both the inner and the outer walls the widths increase to an equilibrium value of 2.9mm and 2.3mm respectively. This shows the expected increase in meltpool volume due to heat retention between passes but also illustrates the ability of the layer height restriction process to cope with excess laser power. A pyrometer is planned for incorporation into the system to provide data on melt pool temperatures.

A pyrometer is planned for incorporation into the system to provide data on melt pool temperatures. This data could be used to reduce laser power when appropriate to compensate for excessive melt pool temperatures.This method of layer height restriction also appears to be useful in building structures with overlapping or crossed layers. The requirement for area surface cladding or solid fills is to overlap parallel clad tracks such that no porosity appears between them. In low dilution cladding it has been stated [11] that the aspect ratio of a clad layer being overlapped in this fashion should not be more than 5 [12] and that it should not overlap the previous layer by more than 70%. If higher dilution is permissible the aspect ratio can be increased in proportion [5].

If higher dilution is permissible the aspect ratio can be increased in proportion [5]. The clad layer height limiting effects of the process described are adequate for controlling the aspect ratio of the clad layer.The following micrographs (Fig 9) show the effect of clad height limiting on partiallyoverlapping layers. Using the same process parameters it has been possible to change the aspect ratio in the clad layers from 5 in the 0.4mm step height case to 3.33 in the 0.6mm case. In each case the overlap has been 50%, the process velocity has been 15mm/s and the powder mass flow rate has been 9.4g/min.

In each case the overlap has been 50%, the process velocity has been 15mm/s and the powder mass flow rate has been 9.4g/min. The laser power was 1 kW with a 2mm spot size using a 190mm focal length lens.4780.4mm step height, 50% overlap0.6mm step height, 50% overlapFig 9 Micrographs of midsection of thick wall built with partially overlapping clad layers (bar size 200 µm )The 0.4mm step height sample shows a coarser microstructure compared to the 0.6mm step height sample that has been the feature of increased clad layer height restriction referred. Dendrites in the 0.4 mm step height sample are thicker and the interlayer banding is more pronounced.

Dendrites in the 0.4 mm step height sample are thicker and the interlayer banding is more pronounced. Pores are not evident in either sample, indicating that dilution was sufficient to avoid their formation.One point with regards to producing thick walls or billets with partially overlapping layers is that the heat retained in local areas of the part is a function of the processing path, therefore careful selection of the overlap pattern is necessary to prevent some areas of the part cooling more slowly than others, with the resultant effects upon microstructure.

In addition, if the part gets hot enough over a large enough area, convective gas currents will disturb the powder cloud geometry of the nozzle, resulting in reduction in clad height limiting effectiveness [9].7.ConclusionsThe layer height restriction process described appears to be able to cope with surface irregularities up to 3.5mm peak-to-trough measurement independently of any height-control feedback system or part-specific programming. This is in addition to preventing cumulative height errors due to process parameter variation or retained heat in the built part. The increased melt pool temperature associated with such restriction, however, can result in increased wetting and/or coarser microstructures. The proposed inclusion of a feedback power control based on melt pool temperature is anticipated to reduce these undesirable effects.8.References [1].

The proposed inclusion of a feedback power control based on melt pool temperature is anticipated to reduce these undesirable effects.8.References [1]. Murphy M, Lee C, Steen W.M, ‘Studies in Rapid Prototyping by Laser Surface Cladding’, Proceedings of ICALEO 1993, pp 882-891 [2]. Capshaw W. ‘Improved performance for motorsports applications through laser deposition technology’, presented at International Conference on Metal Powder Deposition for Rapid Manufacturing, San Antonio April 2002, pp. 69-76 [3]. Taminger K.M.B et al, ‘Solid Freeform Fabrication: An enabling technology for future space missions’ Proceedings of International Conference on Metal Powder Deposition for Rapid Manufacturing, San Antonio April 2002, pp. 51-60. [4]. Mazumder J. et al, ‘Direct materials deposition: designed macro- and microstructure’, Materials Research Innovations 3,1999, publ. Springer-Verlag, pp 118-131.479[5]. Weerasinghe V.M., ‘Laser cladding of flat plates’, Ph.D.

mpensation of Non-Linear Shrinkage of Polymer Materials in Selective Laser SinteringK. Manetsberger, J. Shen, J. Muellers*DaimlerChrysler Research Center, Ulm, Germany * Dornier Research, Friedrichshafen, GermanyAbstractInaccuracies in the selective laser sintering (SLS) process using polymer materials are typically caused by inhomogeneous shrinkage due to inhomogeneous temperature distribution in the powder bed of the SLS machine. These shrinking effects lead to stress in the sintered parts, causing the part to distort. The inhomogeneous shrinkage of benchmark parts has been compensated empirically in a former work. The results cannot been transferred to all geometries, because each geometry requires a specific temperature for laser sintering and, thus, has its own related shrinkage distribution. In a new theoretical approach, shrinkage behavior is to be integrated in a thermal simulation of the SLS process and the thermal shrinkage calculated prior to the building process.

In a new theoretical approach, shrinkage behavior is to be integrated in a thermal simulation of the SLS process and the thermal shrinkage calculated prior to the building process. In the following, experimental data of the temperature- and pressure- dependent shrinkage of laser-sintered powder samples is presented. Possible theories for a physical model of thermal shrinkage are discussed. In particular, these models have to consider granular characteristics such as internal friction, particle sliding, and powder compaction.IntroductionToday rapid prototyping, rapid tooling and rapid manufacturing processes are counted as established means of accelerating the product development cycle in the automobile and aeronautics industries. Thus we see more and more prototype parts, which in the past were traditionally produced using high-cost, time-consuming milling and joining techniques, being replaced by investment cast parts.

Thus we see more and more prototype parts, which in the past were traditionally produced using high-cost, time-consuming milling and joining techniques, being replaced by investment cast parts. And the master forms deployed in investment casting are now typically plastic forms produced by means of selective laser sintering (SLS). In addition SLS processes are used to generate plastic functional prototypes for both the automobile and aeronautics industries. Yet the above applications alone are not what makes SLS techniques so interesting for industry: the growing demand for customized automotive products and the resulting small lots needed for low-volume production has caused a commensurate demand for the fabrication of volume-produced parts using generative processes. As a key manufacturer of premium cars, DaimlerChrysler is driving research in these direct-to-product technologies to accelerate product development and reduce costs.

As a key manufacturer of premium cars, DaimlerChrysler is driving research in these direct-to-product technologies to accelerate product development and reduce costs. Thus, an essential part of our work focuses on the very strict requirements for accuracy in the SLS process to achieve the desired high quality of the parts produced.346Typically inaccuracies in the selective laser sintering process using polymer materials occur due to inhomogeneous shrinkage during the building and cooling processes, which causes stresses and results in distortion of the laser-sintered part. Temperature measurements [1] performed within an SLS machine showed a strong temperature gradient in the x-y plane of a layer, thus making the shrinkage of a geometry dependent on its x and y positions in the build room. Yet part shrinkage does not only depend on the temperature at which the part is subjected to laser sintering, it is also impacted by cooling, i.e.

Yet part shrinkage does not only depend on the temperature at which the part is subjected to laser sintering, it is also impacted by cooling, i.e. the length of time the powder bed retains heat, and by the thickness of the powder layer (causing pressure on the part). Experiments previously performed on sample geometries consisting of hollow rectangular parallelepipeds with equidistant measuring points placed at various x-y locations in the build room showed non-linear shrinkage of the height. We thus follow that a linear scaling of a geometry is not sufficient. A subsequent real geometry scaled non-linear in accordance with the empirical data generated resulted in a greatly improved accuracy of the laser sintering processThe temperature distribution within a build room and the changes in temperature during the building and cooling processes differ with the geometry of each part to be produced. The reason for this lies in the variation in the energy applied by the laser.

The reason for this lies in the variation in the energy applied by the laser. The energy generated by the laser heats the powder and causes the sintering effect, finally dispersing into the powder bed. And due to the very low thermal conductivity of the polymer powder deployed, the build room retains heat over a relatively long period of time. Moreover, shrinkage in geometries which take up a large part of the build area is very different from that occurring in single, smaller parts since shrinkage in large geometries tends to be retarded owing to internal stresses. Hence, the related non-linear scaling function needs to be determined separately for each part to be laser sintered. So, the process has to be run twice: the geometry is built, shrinkage ascertained, the related scaling data generated, and the upscaled geometry compensating shrinkage then built.

So, the process has to be run twice: the geometry is built, shrinkage ascertained, the related scaling data generated, and the upscaled geometry compensating shrinkage then built. To save build time and the related costs, it would therefore seem sensible to calculate shrinkage before laser sintering and update the data model, thus doing away with the need to replicate the building process.The ProcedureExperiments performed with sample geometries have shown that shrinkage occurring in the SLS machine is dependent on time, temperature, and pressure, i.e. the weight of the powder in which the part is embedded. Hence, to be able to calculate and compensate shrinkage, we must first ascertain the values of these three factors.

the weight of the powder in which the part is embedded. Hence, to be able to calculate and compensate shrinkage, we must first ascertain the values of these three factors. The basis for any workable shrinkage compensation must therefore be an SLS process simulation that allows calculation of thermal distribution within the build room during the building and cooling processes.We laid the experimental foundation for the quantification of shrinkage by carrying out dilatometer measurements as a function of time, temperature, and pressure. We measured the changes in height of laser-sintered, cylinder-shaped PMMA samples at temperature and pressure conditions identical to those in the laser-sintering equipment over a period of time of 10 hours. The measured curves were then mathematically set out as fit functions. However, transfer of idealized dilatometer measurements to the real conditions in the SLS machine, e.g.

The measured curves were then mathematically set out as fit functions. However, transfer of idealized dilatometer measurements to the real conditions in the SLS machine, e.g. under consideration of temperature-dependent friction between the part and the bulk powder, will necessitate a more exact examination of the granular characteristics of the material used. Here,347we provide a possible theoretical description of particle sliding under consideration of temperature-dependent cohesion, friction, and sintering.Thermal Simulation as a Basis for Shrinkage CompensationWe decided to compensate shrinkage by means of an FEM simulation which replicates the thermal conditions prevalent within the build room during the building and cooling processes. The heat-conductivity model used for determining heat distribution is described in a previously published work [2].

The heat-conductivity model used for determining heat distribution is described in a previously published work [2]. As shrinkage is a long-term effect, the highly dynamic effects of sintering may be neglected when calculating the heat balance. Exposure of the part and the sintering process itself are instantaneous, i.e. time is not a factor in the transition from the initial density of the powder bed to the final density of the part. This simplification allows us to select suitably large values for both the distance between nodes for the FEM simulation and the time intervals for the FEM calculation, thus greatly minimizing computation time.Initially, we were only able to compute very simple geometries; yet creating an STL interface to the simulation seemed promising for calculation of parts whose geometries are more complex.

As the exposure time for each layer and, thus, the resulting overall time needed for the building process are greatly dependent on the geometry of the part to be sintered, the interface to STL allows specification of the time each layer is subjected to a certain temperature. And the pressure to which each layer is subjected due to the weight of the bulk powder lying on top may easily be ascertained via the current z position of the layer.Figure 1 depicts isothermal distribution inside the build room using the example of a hollow rectangular parallelepiped. For reasons of symmetry, only one-fourth of the build room need be displayed in the figure.Of course, the results of the shrinkage compensation are only as accurate as the thermal simulation itself, i.e. the actual temperature conditions in the build room must be simulated as closely as possible. For this reason we compared the results of the simulation with empirical data.

the actual temperature conditions in the build room must be simulated as closely as possible. For this reason we compared the results of the simulation with empirical data. However, as we were not able to measure the exact temperature within a part, the temperatures within the powder bed near a sintered geometry were compared.Figure 1: Isothermal distribution within the build room for a hollow rectangular parallelepiped.348Sinter wall with holes for placement of the PT100 sensors.xzFigure 2: Comparison of the temperatures measured during experiments and the results of the thermal simulation.Figure 2 compares the measured temperatures with data from the thermal simulation. In the experiment, we inserted PT100 temperature sensors into holes in a sinter wall and measured the temperature in the powder bed. The temperature curve illustrates the cooling of the powder effected by lowering the build platform.

The temperature curve illustrates the cooling of the powder effected by lowering the build platform. Comparison with simulation data showed excellent consistency.Ascertaining Experimental Data for the Shrinkage ModelTo quantify shrinkage of PMMA parts during the selective laser sintering process, we performed dilatometer measurements of the change in height of the samples as a function of time, temperature, and pressure. The resulting data served as the foundation for a physical shrinkage model.Samples and measuring conditions: The samples used in the experiment were cylindrical in shape with a height of 5mm and a diameter of 7mm. At isothermal conditions, the height of the samples was measured over a period of time of 10 hours: temperatures under TG (TG = 123° C) and over TG (Tmax = 160° C) were applied. Pressure on the samples that would occur from the weight of the powder in the SLS machine was simulated during measuring by the application of various weightloads.

Pressure on the samples that would occur from the weight of the powder in the SLS machine was simulated during measuring by the application of various weightloads. A stamp was used to ensure that pressure was distributed equally over the entire surface of the sample. Measurements were performed at four different loads and at no fewer than five different temperatures for each load applied. A very slight bulging of the samples was registered at the highest temperature and the greatest pressure but, as this was negligible, it was disregarded. For all samples, heat-up time was 4 K/min, so that we were able to assume even heating of the small samples. The samples were placed into the dilatometer under two different conditions:3491.) Friction-free shrinkage: the samples did not come into contact with the quartz walls of thedevice.2.) Retarded shrinkage: to be able to estimate the impact of friction between the sample and the powder bed on shrinkage behavior.

As we were unable to achieve selective heating between the powder bed and the sample, polyamide powder was scattered between the sample and the quartz walls of the dilatometer to simulate the different flow and cohesive properties of the PMMA powder and the sample at the different temperatures.Qualitative description of the shrinking curves: Figure 3 illustrates the shrinking curve of a PMMA sample over a period of time of 10h at a temperature of T=140°C and a pressure of p=18.5g/cm2, which corresponds to the weight of a powder bed that is approx. 30cm deep. During the heat-up phase, the sample initially expands due to thermal expansion. Yet, when a certain temperature is reached, shrinkage begins. As thermal expansion is not yet fully complete at this point, the sample is in a mixed state between expansion and shrinking.

Yet, when a certain temperature is reached, shrinkage begins. As thermal expansion is not yet fully complete at this point, the sample is in a mixed state between expansion and shrinking. To more exactly define this state, we subtracted the thermal expansion of the sample from the curve which shows the change in height (red line): this curve thus shows greater shrinkage then the measured curve where thermal expansion is considered.The largest degree of shrinkage (>50%) occurred for all samples within the first 100 minutes with a sharp decrease in the height of the samples registered. After this length of time, shrinking seems to change over to a state of saturation. This is best explained if we see shrinkage as the effect of particles or particle clusters sliding from their original positions to fill empty spaces. This happens when sinter contacts are thermally activated by the temperature applied, which reduces the friction between the individual particles, causing them to slide.

This happens when sinter contacts are thermally activated by the temperature applied, which reduces the friction between the individual particles, causing them to slide. The particles which are the first to slide into vacancies or vacancy clusters are those whose sliding planes are tilted in the direction of the space to be filled out (Figure 4), as this is the path of the least resistance. This explains why the degree of shrinkage is greatest at the beginning of isothermal densification.

This explains why the degree of shrinkage is greatest at the beginning of isothermal densification. The directions of the sliding planes in the samples are distributed statistically and, over time with more and more particles filling vacancies, the concentration of those particles able to slide into vacancy clusters decreases so that shrinkage approaches a saturation value for longer periods of time.Temperature- and pressure-dependency of the final values of shrinkage: As expected, the height of the cylinder-shaped samples after shrinkage can be expressed as a function of the temperature and the pressure applied to the samples. When the final values of shrinkage and the temperature applied are expressed in the form of an Arrehnius plot, it becomes evident that the two values are linear dependent.

When the final values of shrinkage and the temperature applied are expressed in the form of an Arrehnius plot, it becomes evident that the two values are linear dependent. Thus, the dependency between temperature T and shrinkage sfin may be expressed as follows:sfineE cT,with c=const.E may be understood as a thermal energy of activation: the higher the temperature, the greater the number of particles contacts which are thermally activated with the result that a greater number of particles become active, sliding to fill the empty spaces. This increases shrinkage and subsequently the final density.350(Eq. 1)Figure 3: Shrinking curves of a PMMA sample sintered at T=140°C and p=18.5g/cm2. Note that, in the red curve, thermal expansion of the sample during the heating process has been subtracted.

1)Figure 3: Shrinking curves of a PMMA sample sintered at T=140°C and p=18.5g/cm2. Note that, in the red curve, thermal expansion of the sample during the heating process has been subtracted. The temperature curve for the overall process is included.Figure 4: Schematic sketch illustrating the shrinkage mechanism for laser-sintered polymer samples.351Figure 5: Temperature-dependency of the final values ascertained for shrinkage.The dependency between the final values of shrinkage and the pressure applied is also linear (Figure 6). However, increasing mass, i.e. the weight applied, does not have as great an impact on the degree of shrinkage as increasing the temperature. The linear dependency between these two values – shrinkage and pressure – may best be explained as follows. When the contact surfaces between the particles are thermally activated, cohesive strength is reduced.

When the contact surfaces between the particles are thermally activated, cohesive strength is reduced. Similar to Hooke's Law for elastic bodies, pressure causes the particles to be compressed like a spring. Yet unlike a spring, where the compression of the coils is instantaneous, compression of the particles takes much longer. The internal stresses are relieved when the powder particles slide along their planes to fill any available empty spaces. Therefore, after pressure is removed, the samples – again unlike a spring - do not regain their original shape. This process is, in contrast, a relaxation process.Figure 6: Pressure-dependency of the final values ascertained for shrinkage.352Impact of particle friction on shrinkage: To estimate the impact of friction between the bulk powder and the sinter geometry, we carried out the above-described shrinkage experiments using PA powder.

Figure 7 depicts shrinkage measured at three different temperatures: first without friction and then retarded through friction due to particle contact. For all three temperatures, retarded shrinkage is less than friction-free shrinkage by the same absolute value (1%). This behavior would indicate that, within the temperature range considered, contact friction between the bulk PA powder and the PMMA sample is not temperature dependent. At the temperatures selected here, the PA powder is neither cohesive nor flowable enough to significantly increase friction at the contact surfaces. It naturally remains to be examined in future work whether this behavior holds true for PMMA samples laser-sintered in a PMMA powder bed.Integration of shrinking curves in the simulation: To be able to incorporate shrinkage in the thermal simulation, we must first express the change in height of the samples in mathematical form as a function of time, temperature, and pressure.

When the shrinking curves are shown on the log-log scale, it becomes apparent that the curve, for most of its run, follows a line (Figure 8). Hence, according to a power law, the change in height and, thus, the density of the sample, may be expressed as a function of time for longer time periods. Adding the linear dependency of shrinkage s to pressure p renders the following equation:st(aT (cid:215) )cp,with c=const.The exponents of the exponential function and the temperature are linear dependent (Figure 9):aT )(=+ cTc 12,with c1, c2 = const.Figure 7: Comparison of shrinkage of pre-sintered PMMA samples tested first friction-free and then using PA powder to enhance cohesive strength.353(Eq. 2)(Eq.

2)(Eq. 3)Figure 8: Time-dependency of shrinkage on a log-log scale.Figure 9: Temperature-dependency of the power function exponents.Physical Shrinkage ModelsModeling and simulation of shrinkage can be done on two scales: at a microscopic level which tracks an amount of distinct interacting particles, the so-called Distinct Element Model (DEM), and at a macroscopic level. The latter applies special laws for the plastic flow of cohesive granular matter. [3] With respect to the cohesive strength, relevant sintering laws [2] which may354be extended by a pressure dependence need to be found. This can be done, for instance, if the driving surface tension is enhanced by a value proportional to the local pressure and the particle diameter.A simple version for the DEM system is given through an elastic contact model which consists of two different stiffnesses for loading and unloading in order to incorporate plasticity [3]. Also, cohesion strength has to be considered.

Also, cohesion strength has to be considered. The tangential motion of the particles is controlled by a Coulomb-type frictional law. Due to the history- and temperature-dependence of the sintering process, both the cohesion and the friction parameters need not be constant.In principle, this microscopic simulation together with a certain averaging procedure leads to effective macroscopic dynamic laws. In practice, however, the number of particles is limited for computational reasons. Moreover, most current DEM simulations are restricted to two dimensions. Therefore, it is reasonable to try to directly find some macroscopic constitutive equations.fhystrik1rjk2kc dkc dFigure 10: Two-particle contact with overlap d (left), elastic force law with hysteresis and cohesion (right).Several authors have considered macroscopic stress-strain behavior in the context of soil mechanics.

We may start here from a Drucker-Prager formulation of plasticity, which has to be generalized to include cohesion, for instance [3]. However, the resulting model still needs to be extended to make cohesion history- and temperature-dependent. This has not yet been done.For low stresses, i.e. if no external pressing is carried out, a viscoplastic formulation of the constitutive equations would be more appropriate. The related viscosity coefficients will then depend on the density, leading to a non-linear model [4]. Finding the related sintering laws for the polymer powder remains an area for future work.355Summary and Future WorkLaser-sintered polymer parts shrink during the building and cooling processes. Shrinkage, which may be expressed as a function of time, temperature and pressure, is to be computed before sintering and compensated by upscaling the requisite datasets.

Shrinkage, which may be expressed as a function of time, temperature and pressure, is to be computed before sintering and compensated by upscaling the requisite datasets. The basis for shrinkage compensation as set out here is given through a thermal simulation of the SLS process, which replicates the temperature distribution in the build room very accurately. Data from dilatometer examinations of laser-sintered sample geometries serve as the foundation for a shrinkage model. In experiments, the samples demonstrated a change in height which, given longer periods of time, behaves in accordance with a power law with time. The exponents of the exponential function and the temperature proved to be linear dependent. In addition, the final values of shrinkage can be expressed in the form of an Arrehnius plot as a function of temperature and also show a linear dependency to the pressure applied.

In addition, the final values of shrinkage can be expressed in the form of an Arrehnius plot as a function of temperature and also show a linear dependency to the pressure applied. The former function results from the necessary thermal activation of the particle contacts, whereas the latter is explained as a stress-strain relationship similar to Hooke's Law, where internal stresses caused by applying pressure to the particles are relieved by the particles sliding into available vacancy clusters. If particle friction is generated through the addition of bulk powder, shrinkage is reduced significantly. Transfer of the experimental results achieved under the idealized measuring conditions prevalent in the dilatometer to the real conditions in SLS machines will require further work. However, this work offers various theoretical approaches for such transfer.

However, this work offers various theoretical approaches for such transfer. While the resulting shrinkage model will need to consider friction and cohesion between the particles at the microscopic level, relevant macroscopic equations for shrinkage are required to significantly minimize computation time.References1. J. Shen, J. Steinberger, J. Göpfert, R. Gerner, F. Daiber, K. Manetsberger, S. Ferstl, “Inhomogeneous Shrinkage of Polymer Materials in Selective Laser Sintering”, SFF 2000, Austin, Texas2. J. Steinberger, J. Shen. K. Manetsberger, J. Muellers, “Simulation of the SLS Process as the Basis of a Process Optimization”, SFF 2000, Austin, Texas3. Vermeer, P.A, Diebels, S., Ehlers, W., Herrmann, H.J., Luding, S., Ramm, E., University of Stuttgart, Germany, “Continuous and Discontinuous Modelling of Cohesive-Frictional Materials” 2001. XIV, 307 pp. Hardcover4. Riedel.

mpensation zone approach to avoid Z errors in Mask Projection Stereolithography builds Ameya Limaye and Dr. David Rosen George W. Woodruff School of Mechanical Engineering Georgia Institute of Technology Atlanta GA 30332 Contact: Telephone number: 404 894 9668 Email: david.rosen@me.gatech.eduReviewed, accepted August 26, 2005 AbstractPrint-through results in unwanted polymerization occurring beneath a part cured using Mask Projection Stereolithography (MPSLA) and thus creates error in its Z dimension. In this paper, the "Compensation zone approach" is proposed to avoid this error. This approach entails modifying the geometry of the part to be cured. A volume (Compensation zone) is subtracted from underneath the CAD model in order to compensate for the increase in the Z dimension that would occur due to Print-through.

A volume (Compensation zone) is subtracted from underneath the CAD model in order to compensate for the increase in the Z dimension that would occur due to Print-through. Three process variables have been identified: Thickness of Compensation zone, Thickness of every layer and Exposure distribution across every image used to cure a layer. Analytical relations have been formulated between these process variables in order to obtain dimensionally accurate parts. The Compensation zone approach is demonstrated on an example problem.1. IntroductionMask Projection Stereolithography (MPSLA) is an additive fabrication process used to build physical components out of a photopolymer resin. The CAD model of the part to be built is sliced by horizontal planes and the slices are stored as bitmaps. These bitmaps are displayed on a dynamic mask and are imaged onto the photopolymer resin surface.

These bitmaps are displayed on a dynamic mask and are imaged onto the photopolymer resin surface. When a bitmap is imaged onto the resin surface, a layer corresponding to the shape of the bitmap gets cured. This layer is coated with a fresh layer of resin by lowering it inside a vat holding the resin and the next layer is cured on top of it. By curing layers one over the other, the entire part is built. The schematic of the MPSLA system realized at Georgia Tech. is shown in Figure 1.Figure 1 Schematic of the MPSLA system developed at Georgia Institute of Technology326MPSLA can fabricate true 3D geometries as opposed by 21/2D shapes fabricated by other micromachining techniques like silicon etching and LIGA (Madou, 1997) and so, can be a valuable tool to fabricate as well as package MEMS components. The applications like miniature reaction chambers for bio and chemical sensors and micro fluidic devices like micropumps and microvalves have been envisioned in (Tse et al., 2003).

The applications like miniature reaction chambers for bio and chemical sensors and micro fluidic devices like micropumps and microvalves have been envisioned in (Tse et al., 2003). In order to realize this potential of the MPSLA process, the process should be able to fabricate accurate microstructures. The MPSLA process needs to be analytically modeled and efficient process planning method should be formulated to fabricate accurate parts.Since the process is quite new, most work on it is experimental in nature. MPSLA systems have been realized by (Bertsch at al., 1997, Chatwin et al., 1998, Beluze et al., 1999, Chatwin et al., 1999, Farsari et al., 1999, Monneret at al., 1999, Bertsch et al., 2000, Farsari et al., 2000, Bertsch et al., 2001, Monneret et al., 2001, Hadipoespito et al., 2003, Limaye and Rosen, 2004). Most of the research presented in these papers is experimental. In (Limaye and Rosen, 2004) we have presented the MPSLA system realized at Georgia Tech.

Most of the research presented in these papers is experimental. In (Limaye and Rosen, 2004) we have presented the MPSLA system realized at Georgia Tech. The system comprises of broadband UV lamp as the light source, a Digital Micromirror Device (DMD) from Texas Instruments as a dynamic mask and an automated XYZ stage from ASI imaging. We cure parts out of the DSM SOMOS 10120 resin with our system. In (Limaye and Rosen, 2004) we have modeled the lateral dimensions of a layer cured using our MPSLA system in terms of the process parameters. The irradiation of the resin surface has been modeled using the ray tracing approach (Smith, 1990). The curing characteristics of the resin have been empirically modeled by plotting its working curve. In (Limaye and Rosen, 2005), we have used these models to formulate a process-planning method to cure a layer with the required lateral dimensions.

In (Limaye and Rosen, 2005), we have used these models to formulate a process-planning method to cure a layer with the required lateral dimensions. This method is used to generate the bitmap to be displayed on the DMD and compute the time for which it should be imaged onto the resin surface to cure the desired layer. Using this process planning method it is possible to cure layers within 3% error in their lateral dimensions.The vertical dimension of a MPSLA part built by curing dimensionally accurate layers over one another is not equal to the algebraic summation of the individual layer thicknesses and involves some errors. These errors are a result of unwanted curing occurring due to Print through (Section 2). In Section 3, a method called “Compensation zone approach” is proposed to compensate for this unwanted curing. Analytical relations between the various process parameters are derived to avoid the Print through errors using the Compensation zone approach.

Analytical relations between the various process parameters are derived to avoid the Print through errors using the Compensation zone approach. The Compensation Zone approach is applied to an example problem in Section 4.2. Print through errorWhenever a layer is cured, radiation penetrates beyond its intended thickness and ends up exposing the resin underneath it. As the radiation propagates through the resin, it gets attenuated according to the Beer Lambert’s law of absorption. The exposure received by the resin at a depth z from the surface is given by =E e z Dp −/E z ( where E z() ) is the exposure at the depth z in resin,0(1)E 0 is the exposure at the resin surface, Dp is the depth of penetration, which is a resin constant. The resin beneath a MPSLA build gets exposed due to radiation penetrating to it from all the layers cured above it.

The resin beneath a MPSLA build gets exposed due to radiation penetrating to it from all the layers cured above it. Exposure being additive, a point is reached when the exposure received by the resin underneath the MPSLA part equals the threshold exposure for polymerization (Ec).327This unwanted curing causes an increase in the vertical dimension of the part. This error is called as the Print through error. The source of Print through error is depicted pictorially in Figure 2.Commercial Stereolithography machines offer the user an option called “Layer compensation” to compensate for Print through (3D Systems user manual). In this approach, the bottom most layer of a part is skipped to compensate for the increase in vertical dimension that would occur due to Print through. This is an ad-hoc approach, which assumes that the increase in Z dimension due to Print through is always equal to the thickness of the lowermost layer.

This is an ad-hoc approach, which assumes that the increase in Z dimension due to Print through is always equal to the thickness of the lowermost layer. Also, by skipping the lowermost layer, the details at the bottom of the part are lost. Furthermore, the approach can be used to control errors occurring only on horizontal down facing surfaces and is ineffective in controlling the errors occurring in slanting and curved down facing surfaces.Figure 2 Print through error3. Compensation zone approachIn this paper, we propose a method we call as the “Compensation zone approach” to avoid Print through errors in MPSLA builds. This method entails subtracting a tailored volume (Compensation zone) from underneath the CAD model in order to compensate for the increase in the Z dimension that would occur due to Print-through.

This method entails subtracting a tailored volume (Compensation zone) from underneath the CAD model in order to compensate for the increase in the Z dimension that would occur due to Print-through. By controlling the process parameters, including the thickness of the Compensation zone, it is possible, in theory, to eliminate the Print through errors completely.The following are the process variables, i.e.

process parameters under user control.Thickness of the Compensation zone, given by the function Zc(x,y) • Thickness of every layer given by the function LTk(x,y), where LTk(x,y) is the thickness of the kth layer.Exposure supplied to cure every layer, given by function Ek(x,y)328The fixed process parameters are the resin properties:Threshold exposure of polymerization Ec • Depth of penetration Dp, which is a measure of the attenuation of light as it passes through the resin.3.1 Relation between layer thickness LTk(x,y) and exposure Ek(x,y)The exposure Ek(x,y) supplied to cure the kth layer should cure the layer to a depth greater than or equal to the thickness of the layer LTk(x,y) so as to ensure that the layer does not float away. In other words, Ek(x,y) should be such that the exposure at the depth LTk(x,y) should equal to Ec. In order to minimize the Print through error, and thereby the thickness of the Compensation zone, this exposure should be as low as possible.

In order to minimize the Print through error, and thereby the thickness of the Compensation zone, this exposure should be as low as possible. Assuming the Beer Lambert’s law of absorption for radiation attenuation and the threshold model for resin cure (Jacobs, 1992), equation (2) can be derived.E (x, y) E e c≥kLT (x,y)/ D pk(2)3.2 Relations between layer thickness LTk(x,y) and Compensation zone Zc(x,y)When the kth layer is cured, radiation would penetrate beyond its intended thickness and expose the bottom surface of the part.

Assuming that the radiation is attenuated according to the Beer Lambert’s law, the exposure received by the bottom surface due to radiation coming from the kth layer will be given by equation (3).kE (x, y) E k=bk(, x y) e∑− (m1 =LT (x,y) m+, ( Z x yc))/ Dp(3)where the exponential termk ∑ 1 m =LT (x, y) m+( Z x y c,)is the height of the top of the kth layer from thebottom surface.The bottom surface would receive exposure penetrating to it from all the n layers curedabove it.

So, the total exposure received by the bottom surface will be given by( E x y b,) =n∑ E (x, y)bk(4)k=1where n is the number of layers the part is made up of.Substituting the value of E (x, y) is given by equation (5).bkfrom equation (3), the exposure received by the bottom surfacek( E x y b,)=n∑ Ek(, x ye )∑− (m1 =LT (x,y) m+, ( Z x yc))/ Dp(5)k=1329The exposure received by the bottom surface should be equal to Ec in order to haveaccurate vertical dimension, i.e.( E x y b,) =E c(6)From equations (5) and (6),kn∑E k(, x y) e−(∑m1 =LT (x,y) m+, ( Z x yc))/ Dp=E c(7)k=1Substituting the value of Ek x y(,) from equation (2) into equation (7),kn∑E e cLT (x,y) D p/ke−(∑m=1LT (x,y) m+( , Z x yc))/ D p=cE(8)k=1Canceling the term Ec in equation (8), we get the first relation between the layer thickness LTk(x,y) and Zone of compensation Zc(x,y) as given by equation (9).kn∑eLT (x,y) D p/ke−(∑m=1LT (x,y) m+( , Z x yc))/ D p=1(9)k=1The height of the part is the algebraic summation of the thickness of every layer and the thickness of the Compensation Zone.

The height of the part can be readily obtained from its CAD model as h(x,y), and will be given by equation (10). This is the second relation between the layer thickness LTk(x,y) and Zone of compensation Zc(x,y)n, h x y ()=∑( LT x y k,)+( Z x y c,)(10)k=1In order to build a MPSLA part with accurate vertical dimensions, process parameters should be selected so that equations (2), (9) and (10) are satisfied.

The Compensation Zone approach can be represented as a problem of solving simultaneous equations as follows.330GivenGeometry of the part h(x,y) • Resin properties: Depth of penetration Dp and Threshold exposure of polymerization EcFindThickness of all layers LTk(x,y) • Thickness of Zone of Compensation Zc(x,y) • Exposure supplied to cure every layer Ek(x,y)SatisfyE (x, y) E e c kE (x, y) E e c k(Equation (2) in Section 3.1)kn∑e( ( LT (x,y) k+OC x y ( , k))/ D ) pe(−∑m1 =LT (x,y) m+Z x y , (c)/ D ) p=1(Equation 9 in Section 3.2)k=1( • h x y,)=n∑( LT x y k,)+( Z x y c,)(Equation 10 in Section 3.2)k=1The above problem has more variables than equations and is thus an under-constrained problem with multiple solutions. So, additional constraints, like those on vertical resolution, build time etc. can be imposed on the process planning.4.

So, additional constraints, like those on vertical resolution, build time etc. can be imposed on the process planning.4. Implementing the Compensation Zone approachThe problem formulated in Section 3 involves more variables than equations. So, the simultaneous equations would have multiple solutions. In order to have a unique solution, additional constraints would have to be imposed on the process. The solution can then be computed using a root finding algorithm. In this section, the approach is explained by considering an example.Example: The micro part with dimensions as shown in Figure 3 is to be built in a resin with known values of Ec and Dp. The height of the part is 500µm. The part has a horizontal and a slanting down-facing surface. The slanting down-facing surface is at an angle 45° to the horizontal plane.

The height of the part is 500µm. The part has a horizontal and a slanting down-facing surface. The slanting down-facing surface is at an angle 45° to the horizontal plane. Determine the thickness of every layer, the exposure to be supplied to cure each of these layers and the thickness of the Compensation Zone that has to be subtracted from the beneath the part, in order to cure the down facing surfaces accurately.331Figure 3 MPSLA part built to validate the Compensation Zone approachSolution:The following simplifying assumptions are made in order to solve the problem.The thickness of every layer is constant and equal to 100µm except near the edges of the part. By this assumption, only the layer just above the compensation zone will have a thickness other than 100µm, while all other layers will have a thickness equal to 100µm.

By this assumption, only the layer just above the compensation zone will have a thickness other than 100µm, while all other layers will have a thickness equal to 100µm. Mathematically, this can be represented as For any point (x1,y1), 0<LTt(x1,y1)<100, then LTk (x1,y1) = 0 or LTk (x1,y1) = 100, for all k from 1 to n but k not= t LTm(x1,y1) ≥ LTn(x1,y1) for all m>n(11)This thickness of the Compensation zone is to be kept to its minimum. For this, each layer would have to be supplied with the minimum possible exposure. For any point (x1,y1), E (x , y ) E e (12) = 1 c With these two assumptions, the problem now has a unique solution.

The value of Zc(x1,y1) which would converge the functionkn∑e(LT (x ,y )/ D ) pk11e(−∑m=1LT (x ,y ) 1 1m+, Z x y(c11)/ D ) pk=1to its desired value, i.e. 1, can be determined by binary search. To start the binary search, the minimum value and maximum values of Zc(x1,y1) are assumed to be 0µm and 100µm respectively. Once the values of Zc(x1,y1) and LTk(x1,y1) are determined by binary search, the values of Ek(x1,y1) can be determined using the fourth constraint.The XY extents of the example part are 900µmX500µm. These extents were discretized into 18X10 grid of points. At each of the points, the values of Zc and LTk and Ek were determined. The grid was kept this sparse to reduce the computation time. In theory, the separation between grid points can be made as small as the size of a pixel irradiated by a single micromirror on the mask.

The grid was kept this sparse to reduce the computation time. In theory, the separation between grid points can be made as small as the size of a pixel irradiated by a single micromirror on the mask. 18X10 matrices for the Zone of compensation, the layer thicknesses and the exposures to be supplied to the layers have been generated. The profile of the part that would be cured has been simulated. As shown in Figure 5, the Compensation zone to be subtracted at the horizontal down facing surface of the part is 14.7µm. The Compensation zone to be subtracted from underneath a layer varies depending upon the distance of the layer cured on top of it. For example, the Compensation zones denoted A and B in Figure 5 receive the same exposure from layer 3. However the layers 4 and 5 are situated closer to Compensation zone A than B and so, A receives more radiation penetrating from them than does B.5.

However the layers 4 and 5 are situated closer to Compensation zone A than B and so, A receives more radiation penetrating from them than does B.5. ConclusionsThe Compensation Zone approach ensures that the bottom most surface of the part being built receives an exposure exactly equal to the threshold exposure of polymerization (Ec). This approach however is based upon certain assumptions. It assumes the threshold model of exposure, i.e. there is no partial curing, it assumes that the attenuation of radiation as it passes through liquid resin and cured layers are the same. Further, the approach also assumes that the exposure is additive. Due to these assumptions, the Compensation Zone approach is unlikely to eliminate the Print through error completely.The Compensation Zone approach can be adopted to reduce the problem of stair stepping to create layers conforming to slanting and curved part boundaries. For example, in the Figure 4, the stair steps are of size 50µm.

For example, in the Figure 4, the stair steps are of size 50µm. The size of the stair-steps can be further reduced by increasing the density of the matrix used to populate the part near its edges. This can improve the surface finish on the down facing surface of the part further.334ABFigure 5 Simulated profile of the part that would be built using the Compensation zone approach6. Future WorkThe Compensation Zone approach shall be used to build the example part on the MPSLA machine at Georgia Tech whose dimensions shall be compared with those of the simulated part. The MPSLA process needs to be made robust against changes in resin properties Ec and Dp caused by unwanted radiation (e.g., ambient light, print-through) that can cause partial polymerization.

The MPSLA process needs to be made robust against changes in resin properties Ec and Dp caused by unwanted radiation (e.g., ambient light, print-through) that can cause partial polymerization. A method to compute the values of the process variables so that the process becomes robust against change in resin properties shall be formulated.335AcknowledgementsWe gratefully acknowledge the financial support from the RPMI member companies, the Manufacturing Research Center at Georgia Tech., the Georgia Research Alliance (grant GRA.IC05.F), and the National Collegiate Inventors and Innovators Alliance (NCIIA).ReferencesBeluze, L., Bertsch, A., Renaud, P. 1999. Microstereolithography: a new process to build complex 3D objects. SPIE Symposium on design, test and microfabrication of MEMS/MOEMS 3680(2): 808-17.Bertsch, A., Zissi, S., Jezequel, J., Corbel, S., Andre, J. 1997. Microstereolithography using liquid crystal display as dynamic mask-generator.

Materials made by the laser powder bed fusion (L-PBF) process contain unavoidable process-induced porosity whose effect on the mechanical properties needs to be considered for fatigue applications. Results from this work show that not all pores in L-PBF parts promote fatigue crack initiation. The length scale of local microstructure defects, i.e. grain boundary, could be larger than the pores and in such cases they are the primary cause for crack initiation. Samples were produced in this work to demonstrate the critical defect size responsible for the transition between the porosity-driven and microstructure-driven failure modes.Keywords: Fatigue; porosity; microstructure; powder bed fusion; selective laser melting; stainless steel 316L1. IntroductionAdditive manufacturing (AM) is a family of advanced manufacturing methods which produces near net-shape parts with complex geometry based on input from computer-aided design.

IntroductionAdditive manufacturing (AM) is a family of advanced manufacturing methods which produces near net-shape parts with complex geometry based on input from computer-aided design. Its forming principle is different from conventional subtractive manufacturing methods, as objects are built in a layerwise manner. Thus, while offering substantial design freedom and more efficient utilization of resources, different mechanical properties and applications are expected of parts built by AM processes. Laser powder bed fusion (L-PBF) is a popular metal AM process which uses high power laser as the energy source to melt powdered raw material. It has been widely explored as the means of manufacturing for the aerospace, automotive and medical industries for producing lightweight and novel structures.Stainless steel 316L is an extra-low carbon austenitic chromium-nickel stainless steel which has shown good compatibility with the L-PBF process.

It is frequently used for its good corrosion resistance, particularly in biomedical and marine-related applications. Fully dense stainless steel 316L parts with superior strength to that of the conventional manufacturing processes and good ductility have been produced [1]. However, concerns regarding the fatigue behaviour of L-PBF parts had been raised as process-induced porosities were shown to cause365fatigue crack initiation and premature failure [2-7]. As the applications of stainless steel 316L often desire good fatigue resistance, such as in the case of orthopaedic implants, the influence of L-PBF process parameters and the resulting porosity level on this particular aspect of mechanical property requires some due attention.It is known that fatigue failure occurs as a result of the competition between microstructure and voids in polycrystalline materials [8, 9].

Conventional manufacturing processes, such as casting and welding, produce parts containing microcavities and shrinkage defects. For L-PBF, the localized heating and rapid cooling, coupled with the use of powder as the feedstock, can lead to the formation of defects, such as lack of fusion and gas pores. The choice of process parameter settings, such as layer thickness, laser power, scan strategy and hatch, strongly influence the shape, size and number of the defects [10]. In comparison with a number of alloys such as Ti6Al4V, Inconel 718 and carbon steel, stainless steel 316L was shown to be one of the alloys with the highest susceptibility to forming lack of fusion defects [11]. This is possibly a results of its high thermal conductivity, which reduces its ability at retaining heat [12]. Any process inconsistency, e.g.

This is possibly a results of its high thermal conductivity, which reduces its ability at retaining heat [12]. Any process inconsistency, e.g. uneven powder layers being deposited due to damaged recoater and process parameter deviations from the preset values, may subject stainless steel 316L to defect formation even under optimized processing conditions. This implies that stainless steel 316L components produced by L-PBF could be prone to defect-induced failure under cyclic loading.When acceptable control of defect size is achieved, e.g. by the choice of appropriate processing parameter, such that the defects are small in comparison to the local microstructure heterogeneities, they may be considered as benign. Conventionally, non-metallic inclusions smaller than 5 μm are usually harmless for low purity parts [13]. Precise determination of the critical pore size is difficult due to the random arrangement of local microstructure.

Precise determination of the critical pore size is difficult due to the random arrangement of local microstructure. For example, the non-propagating crack size of a 0.13% C steel was found to be in the order of 100 μm, even though the average grain size was only 34 μm [14]. For L-PBF stainless steel 316L, studies on this aspect of the fatigue behaviour is not available. In terms of the influence of microstructure on fatigue property, a number of studies had highlighted the importance of grain size and texture on fatigue crack initiation and growth. Leuders et al. [15] and Riemer et al. [16] tested the effect of stress relieving heat treatment and hot isostatic pressing on the high cycle fatigue and crack growth behaviour of stainless steel 316L respectively. Both concluded that residual stress and porosity have insignificant impact on the fatigue resistance and attributed such behaviour to its high ductility.

Both concluded that residual stress and porosity have insignificant impact on the fatigue resistance and attributed such behaviour to its high ductility. As the austenitic structure of stainless steel 316L does not undergo phase transformation during cooling, it does not contain the martensite phase and therefore brittle failure is avoided. This could have led to the high ductility and defect-tolerant behaviour [16]. Another source of microstructure defects in L-PBF parts is the melt pool boundary. It was found to impact failure under monotonic tensile loading and give rise to anisotropic properties [17]. Anisotropic fatigue behaviour had also been observed in L-PBF-processed stainless steel 316L parts, where the vertically-built specimens produced shorter fatigue life [3, 18].

Anisotropic fatigue behaviour had also been observed in L-PBF-processed stainless steel 316L parts, where the vertically-built specimens produced shorter fatigue life [3, 18]. The competing influence of grain, porosity and melt pool boundary could be responsible for such behaviour and this needs to be addressed in detail.As existing studies did not provide explicit information on the density and porosity conditions of their samples, quantitative correlations between porosity, microstructure and fatigue properties cannot be made from the previous works. Considering that there is a lack of international standards for the certification and quality assurance of part made by powder bed366fusion, a systematic examination of the process-structure-property relationship for L-PBF is desirable for the adoption of this process for structural applications.

In view of this gap, this study aims to provide qualitative and quantitative evaluation of the competing influence of porosity and microstructure on the high cycle fatigue properties of L-PBF stainless steel 316L. Experimental and analytical results from this work can be used as a preliminary guide on the possible failure modes for L-PBF parts with different porosity conditions.2. Experimental SetupCommercially available stainless steel 316L powder with spherical shape and mean diameter of 17 μm was used. Four sets of samples were fabricated by the EOS M 290 system at layer thickness settings of 20 μm, 40 μm, 60 μm and 80 μm, while keeping all the other process parameters, such as laser power, scan speed and hatch, constant; they are referred to as T1, T2, T3 and T4 respectively in this work.

Rectangular blocks were built with the long edge being oriented at an angle from the recoating direction, as shown in Figure 1(a), to prevent uneven powder deposition due to the interaction between the part and the recoater blade [19]. After being removed from the baseplate, the blocks were subsequently machined to produce the samples for mechanical testing. Small test blocks were built adjacent to them for porosity characterization and metallographic examination.Density of the test blocks was measured by the Archimedes method (Mettler Toledo XS204) and micrograph analysis of planes parallel to the build direction. Since fatigue failure is sensitive to the largest occurring defects, inspection was also performed to estimate the size distribution of the large defects. This was done by taking 30 random images of areas 0.31 mm2 each using optical microscope and identifying the largest defect size within each image [20, 21].

This was done by taking 30 random images of areas 0.31 mm2 each using optical microscope and identifying the largest defect size within each image [20, 21]. The equivalent spherical pore diameters were then calculated and used as the input for fitting the Gumbel Extreme Value Distribution, following the maximum likelihood method as described in the ASTM E2283. The cumulative distribution is given by𝐹(𝑥) = exp⁡(− exp (−𝑥−𝜆𝛿)) (1)where x is the equivalent spherical pore diameter, λ and δ are the location and scale parameters of the Gumbel distribution respectively.Fatigue test samples were designed in accordance to ASTM E466 (Figure 1 (b)). They were machined from the rectangular blocks in Figure 1 (a) by the wire cut electrical discharge machining (EDM) method, followed by manual grinding to remove the resulting surface layers. Samples were extracted from each block along the z-axis such that the loading direction was perpendicular to the build direction.

Samples were extracted from each block along the z-axis such that the loading direction was perpendicular to the build direction. Fatigue tests were conducted in air under a constant force- controlled sinusoidal loading at R=0.1 on a servo hydraulic testing system (MTS 810) at a frequency of 5 Hz.367Figure 1 (a) Rectangular blocks and small test blocks were built at an angle from the recoating direction, (b) geometry of fatigue test samples (dimensions in mm). Eight samples were extracted from each rectangular block along the z-axis by wire cutting.3. Results3.1 Porosity – Volume Fraction and Size DistributionOptical micrographs in Figure 2 (a)-(d) show the porosity on the vertical planes of the test blocks. T1 and T2 contain predominantly small spherical pores. For T3, there is a sudden increase in the pore size and the pores are irregular in shape.

T1 and T2 contain predominantly small spherical pores. For T3, there is a sudden increase in the pore size and the pores are irregular in shape. This indicates that sufficient remelting of the previously deposited powder layer can no longer be achieved at this layer thickness, resulting in the formation of lack of fusion defects. Further increase in layer thickness at T4 led to a large increase in the number of lack of fusion defects, such that they were in closer proximity to each other. This observation is consistent with the measured relative densities shown in Table 1, except that the Archimedes method is unable to capture the small change in porosity between T1 and T2.Figure 2 (a) – (d) Optical micrographs showing porosity (in black) on planes parallel to thebuild direction for T1, T2, T3 and T4 respectively.368The Gumbel distribution parameters of the samples are also shown in Table 1. The value of λ is equal to the mode of the Gumbel distribution, i.e.

The value of λ is equal to the mode of the Gumbel distribution, i.e. the equivalent spherical pore diameter with the highest frequency of occurrence in a sample. For near fully dense parts, i.e. T1 and T2, most of the pores are in the order of 10 μm. At close to 99% relative density (based on the image analysis method), i.e. T3, large pores in the order of 50 μm are already present. Also, in addition to the increase in pore size, the pores are more widely distributed with increasing layer thickness, as shown by the values of the scale parameter δ. This is mainly due to stereological effect where cutting sample for metallographic examination exposed the irregularly-shaped defects at different sections along their length. In fact, for T3 and T4, the assumption of spherical defects in estimating the pore size is no longer valid. The orientation of the pores with respect to the loading direction is important for estimating the size of these defects.

The orientation of the pores with respect to the loading direction is important for estimating the size of these defects. However, for simplicity, spherical defects were assumed in this work.

However, for simplicity, spherical defects were assumed in this work. Lastly, in comparison with the Archimedes method, the Extreme Value Distribution was able to capture the small increase in pore size from T1 to T2, as indicated by the larger λ value.Table 1 Density of the samples measured by the Archimedes method and image analysis of the optical micrographs, and the Gumbel Extreme Value Distribution parameters for describing the equivalent spherical pore diameter distributions of the largest-occurring defects.Sample IDT1 T2 T3 T4Layer thickness (μm)20 40 60 80Relative density (%)ArchimedesImage analysis98.93±0.08 99.14±0.01 98.11±0.02 91.95±0.1499.98±0.05 99.92±0.18 98.88±0.95 93.66±3.15Gumbel distribution parameterLocation parameter λ (μm) 6.6 13.4 53.6 146.6Scale parameter δ (μm) 4.3 7.6 35.9 51.73.2 High Cycle Fatigue and Facture BehaviourFatigue LifeFigure 3 shows the fatigue life of the samples. T1 and T2 are nearly identical.

T1 and T2 are nearly identical. Using T1 as the reference, the fatigue lives of T3 and T4 were reduced significantly, by 65% and 94% respectively. The change in fatigue resistance is proportional to the relative densities of the samples, indicating that porosity plays an important role on the failure of T3 and T4 relative to T1 and T2.369Figure 3 Fatigue life of laser powder bed fusion stainless steel 316L samples subjected tomaximum cyclic loading of 438 MPa.FractographyFigure 4 displays the SEM images of the fracture surfaces after fatigue tests. The dotted lines in Figure 4 (a) and (c) separate the fatigue crack origins from the crack propagation region for T1 and T2 respectively. Figure 4 (b) and (d) are the enlarged views of the crack initiation regions. Both samples exhibited intergranular crack initiation behaviour. The arrows in Figure 4 (b) and (c) indicate intergranular de-bonding at grain boundaries.

Both samples exhibited intergranular crack initiation behaviour. The arrows in Figure 4 (b) and (c) indicate intergranular de-bonding at grain boundaries. Comparing the two samples, T2 contained a larger crack initiation region with more extensive intergranular cracking. Figure 4 (d) shows a large area of neatly arranged dendritic cells. The presence of intergranular fracturing and micro-cracks is explained as follows: during solidification, the rapid cooling process could have resulted in competitive grain growth at the solid-liquid interface and caused the formation of differently oriented grains at the melt pool [22] and a large amount of dislocations, even though the global grain orientation is parallel to the build direction, i.e. epitaxial grain. Material deformation by the slip mechanism is constrained by the grain boundaries, where dislocation pile-up can trigger intergranular fatigue cracking [23].

epitaxial grain. Material deformation by the slip mechanism is constrained by the grain boundaries, where dislocation pile-up can trigger intergranular fatigue cracking [23]. Increasing the layer thickness could have led to less extensive remelting and grain regrowth. Grain boundary defects were thus more prevalent for T2 than T1, contributing to the larger crack initiation region.SEM image of the fracture surface of T3 in Figure 4 (e) shows fatigue crack initiation from a surface defect in the order of 100 μm. This is comparable with the defect size obtained from the Extreme Value Distribution. Materials transited rapidly into the propagation mode once cracks were initiated. Cracks could have been initiated at many other lack of fusion defects in the material subsurface. They joined the main crack propagation path and became visible on the fracture surface. Overall, these lack of fusion defects had resulted in the near 60% reduction in fatigue life for T3 as compared to T1.

Overall, these lack of fusion defects had resulted in the near 60% reduction in fatigue life for T3 as compared to T1. For T4, the influence of porosity on fatigue failure became even more significant. No distinct crack origin could be identified in Figure 4 (f). The large amount of defects significantly reduced the load-bearing capacity of the material. The interaction of closely spaced pores compounded the effect by increasing the local stress concentration and multiple cracks were initiated simultaneously, leading to the much lower fatigue strength.370Figure 4 SEM images of sample fracture surfaces.

(a) T1 showing transition from mixed intergranular and transgranular fracture at the crack initiation site (within dotted line) to transgranular propagation, (b) enlarged view of (a) showing debonding at sub-grain boundaries (white arrows), (c) T2 showing a larger and more intergranular crack initiation region than T1, (d) enlarged view of (c) showing sub- grains revealed by intergranular fracture, (e) T3 showing crack initiation from lack of fusion defect at the sample surface, and (f) T4 showing extensive lack of fusion defects and no discernible major crack origin.3714. Discussions – Competition between Microstructure and PorosityBased on the experimental fatigue test results, it appears that the failure mechanism of T1 and T2 is different from T3 and T4. Specifically, fatigue crack initiation in T1 and T2 was predominantly driven by grain boundary defects, while T3 and T4 failed due to crack initiation at porosity.

Specifically, fatigue crack initiation in T1 and T2 was predominantly driven by grain boundary defects, while T3 and T4 failed due to crack initiation at porosity. This section provides the discussions on the effect of microstructure and porosity on the observed properties.4.1 Fatigue LimitFor low cyclic stress conditions, the size of the local plastic zone is in the order of short range microstructure and the influence of local microstructural heterogeneity is very significant [24]. Consequently, the growth of small and large fatigue cracks are fundamentally different at applied cyclic stress close to the fatigue limit. The Kitagawa-Takahashi concept had been used for defining the critical defect size a0 below which fatigue strength is not affected by changes in crack size [9]. Based on the work by Kitagawa and Takahashi [25], the relationship between crack length a and the stress range for crack propagation ΔS can be described by the equation derived by El Haddad et al.

Based on the work by Kitagawa and Takahashi [25], the relationship between crack length a and the stress range for crack propagation ΔS can be described by the equation derived by El Haddad et al. [26]:Δ𝑆 =Δ𝐾th 𝐹√𝜋(𝑎+𝑎0)(2)where ΔKth is the threshold stress intensity factor range, F is a geometric factor and a is the crack size. For a defect-free component with fatigue limit ΔSe, a0 is a material constant described by:𝑎0 = (Δ𝐾th 𝐹Δ𝑆e)2 1 𝜋(3)A planar near surface circular defect was modelled to evaluate the influence of defect size on the fatigue crack propagation stress, i.e. the fatigue limit. Experimentally determined fatigue crack growth data [16] and the BS7910 [27] for calculating F were used. Figure 5 shows the Kitagawa-Takahashi diagram in terms of Equation (2) and the fatigue limit results obtained in this work.

Figure 5 shows the Kitagawa-Takahashi diagram in terms of Equation (2) and the fatigue limit results obtained in this work. Since λ is equal to the mode of the Gumbel distribution, its values, as presented in Table 1, are representative of the equivalent spherical pore diameters of the samples and were used for estimating the crack size a. As two cracks could be considered as originating from the two ends of the planar defect, the crack size a = λ/2. It can be seen from Figure 5 that cracks originating from process-induced pores for T1 and T2 are in the short fatigue crack regime where fatigue limits are not influenced by the crack size. This agrees with results from the fractography analysis, where pores in T1 and T2 were too small that cracks originated at the grain boundary defects instead. Porosity-induced cracks in T3 and T4 are approaching the transition region between short and long fatigue cracks.

Porosity-induced cracks in T3 and T4 are approaching the transition region between short and long fatigue cracks. For T4, the model considerably over-estimates the fatigue limit; this is because rather than containing an isolated pore, as assumed by the model, sub- in L-PBF produces parts with homogenously distributed defects. optimal processing Neighbouring pores were in close proximity to each other and imposed greater stress concentration, resulting in the lower fatigue strength. Also, the assumption of spherical defects in deriving λ could have underestimated the actual pore size of T3 and T4, where irregular pores were prevalent. An appropriate factor on strength needs to be applied to account for these effects. In summary, process-induced pores in T1 and T2 were in the short fatigue crack regime and were too small to cause crack initiation. Crack initiation from grain boundary led to the similar fatigue372limits of these samples.

Crack initiation from grain boundary led to the similar fatigue372limits of these samples. For T3 and T4, process-induced pores were large enough to cause fatigue crack initiation, leading to the crack size-dependent fatigue limits.Figure 5 Kitagawa-Takahashi diagram comparing the El Haddad model and the experimentalhigh cycle fatigue test data from this work.4.2 Critical Defect SizeIt is of practical interest to define the critical pore size for ensuring the reliability of components under high cycle fatigue loading. The Kitagawa-Takahashi analysis can only be used to provide theoretical but not accurate estimate of the critical pore size. Experimental results from this work shows that this transition exists between the pore size of T2 and T3, such that 13.4 μm < λcritical < 53.6 μm. This should be interpreted in conjunction with the size of the local microstructure.

This should be interpreted in conjunction with the size of the local microstructure. Figure 7 shows the microstructure on the plane parallel to the build direction for a typical L-PBF stainless steel 316L. As observed from the fractographs, intergranular failure could have occurred at the boundary between differently oriented grains. The lengths at such grain boundaries are marked in Figure 6. They are mostly in the order of 30 μm to 40 μm. Since the pores in T2 are smaller, microstructure boundaries become the detrimental defects. For T3, the pores are larger than the size of the local microstructure and they become the most critical defects that trigger fatigue cracking.

For T3, the pores are larger than the size of the local microstructure and they become the most critical defects that trigger fatigue cracking. However, it is to be noted that while results from this work provided evidence to the critical defect size in L-PBF stainless steel 316L, they may be limited to the processing parameter region (optimised for a layer thickness of 20 μm) and scan strategy used in this work, as changing the process settings could lead to different microstructure length scales.373Figure 6 SEM image showing characteristic microstructures of L-PBF parts. The lengths of theboundaries between some different oriented grains are shown.5. ConclusionsThe following conclusions were drawn from the results obtained in this work:Two fatigue failure modes are possible for stainless steel 316L manufactured by the L-PBF process depending on the level of porosity.

For parts containing small pores, grain boundary defect-driven crack initiation is the predominant failure mode (T1 and T2). Parts containing large pores are prone to porosity-driven crack initiation (T3 and T4), which significantly reduced the high cycle fatigue life. When pores are too large and close to each other (T4), local stress concentration increases due to interaction among the pores, resulting in much greater reduction in fatigue strength.1)Having established that the transition in fatigue crack initiation modes occurred between T2 and T3, their porosity parameters were compared with the size of the local microstructure heterogeneity. The pore size and volume fraction of T2 could be used as a conservative estimate of the critical pore size that triggers porosity-induced fatigue crack initiation.2)References[1] I. Tolosa, F. Garciandía, F. Zubiri, F. Zapirain, A.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCOMPLEX FUNCTIONAL SURFACE DESIGN FOR ADDITIVE MANUFACTURINGS. J. Morris1, J. P. R. Dudman1, L. Körner1, P. Melo1, L. H. Newton1, A. T. Clare21 Additive Manufacturing & 3D Printing Research Group, Uni. of Nottingham, NG7 2RD, U.K. 2 Advanced Manufacturing Technology Research Group, Uni. of Nottingham, NG7 2RD, U.K. e-mail: ezxsjm@nottingham.ac.uk (S. J. Morris)ABSTRACTThis paper presents a new methodology for the creation of advanced surfaces which can be produced by Additive Manufacturing (AM) methods. Since there is no cost for enhanced complexity, AM allows for new capabilities in surface design. Micro-scale surface features with varying size, shape and pitch can be manufactured by Two-Photon Polymerisation (2PP).

Micro-scale surface features with varying size, shape and pitch can be manufactured by Two-Photon Polymerisation (2PP). Computer-Aided Design (CAD) tools allowing for this variation to be incorporated into the surface design are only just emerging. With the presented methodology, surfaces are created from a single feature design. Variation is applied to the surface features through algorithmic design tools, allowing for arrays of hundreds of unique features can be created by non-CAD experts. The translation of these algorithmic expressions from CAD to a physical surface is investigated. Using the proposed methodology, 2PP is used to create quasi stochastic surfaces for the purpose of enhancing the biointegration of medical implants against current state-of- the-art.1. INTRODUCTIONModification of surface topology enables products to exhibit enhanced functionality [1]. Nature has long been a source of inspiration for tackling engineering challenges [2]. In 2003, Geim et al.

Nature has long been a source of inspiration for tackling engineering challenges [2]. In 2003, Geim et al. [3] produced samples of micro-fabricated polyimide hairs to mimic keratin hairs found on gecko feet. These exhibited enhanced tactile properties for adhesion. In more recent research, Wen et al. [4] replicated shark denticles to reduce drag force. A single denticle design was patterned onto a flexible surface, resulting in a 6.6% increase in swim speed compared to smooth surfaces. In both studies, the features were of only one design type. In other words a common unit cell of fixed X, Y, Z dimension was created and then arranged in a uniform array with fixed X, Y pitches. This is illustrated in Figure 1 (a). Actual structures in nature exhibit natural variations in these parameters: for example shark denticles show a considerable difference in shape and pitch across the animals’ body as well as across species [4].

Incorporating such variation into a surface array, as depicted in Figure 1 (b), could potentially improve biomimicry and further increase part performance.2423(a)(b)ZYY pitchXX pitchFigure 1: Diagrams illustrating parameters that define an array of surface features. (a) Surface features of the same X, Y, Z dimensions with the same X, Y pitches such as those created in [3-4]. (b) Surface features with varying X, Y, Z dimensions and X, Y pitches.Regenerative medicine could significantly benefit from improved biomimetic surface design, allowing tissue scaffolds to better replicate the surrounding biological material when implanted. This could facilitate the regeneration of damaged or dysfunctional tissue. The interaction between cells and an implanted scaffold depends on surface topography amongst other factors, including surface chemistry.

The interaction between cells and an implanted scaffold depends on surface topography amongst other factors, including surface chemistry. Studies determining cell response to simple features such as pits, grooves and pillars have yielded varying results, depending on: feature size, cell type and the manufacturing method used [5]. Some observed effects include increased cell adhesion, cell differentiation, directed growth and enhanced rates of spreading. Incorporation of nanoscale disorder into the spatial arrangement of surface features can lead to improved cell responses compared to ordered arrays such as: increased cellular adhesion and bone mineral production within mesenchymal stromal cell populations [6-7]. The ability to control the surface topology would allow for customised functionality of tissue scaffolds.To build a surface with parametric variation requires a manufacturing technique that is able to produce these complex designs.

Techniques such as transfer printing and laser ablation have been used to modify surfaces [8-9]. Whilst sub-100nm resolutions have been reported with transfer printing, the process requires the creation of a stamp for each new design. Both techniques are also limited in the possible structures that can be built due to being stamping and subtractive methods of processing. Additive Manufacturing (AM) does not require moulds or templates, building parts directly from CAD files [10]. Moreover AM grants greater geometric freedom, allowing for the creation of complex surface features unobtainable by traditional manufacturing techniques. This makes AM ideal for the production of surface features with a wide variation in morphology and spacing.Current CAD tools are not tailored to designing surfaces with large numbers of non- identical surface features; simple patterning tools can replicate a feature and introduce variance from data defined in an external spreadsheet [11].

However this does not allow for synchronous design: where the CAD model is updated at the same time the user changes an input parameter. Currently users are required to know exactly what parameter values they desire beforehand, manually input these into the dataset, and wait for the model to update. Synchronous design,2424gives the user visual feedback whilst designing, enabling a more intuitive and organic design process. Grasshopper® is an algorithmic modelling, visual programming plugin for Rhinoceros® CAD software. Parameters can be defined by mathematical expressions, allowing the geometric design of a large quantity of features to be changes simultaneously. The program allows for synchronous design and scripts with input controllers (such as sliders) to be written. This makes the design process more accessible to CAD novices.In this paper, a methodology is proposed for designing complex surfaces that can be manufactured by AM.

This makes the design process more accessible to CAD novices.In this paper, a methodology is proposed for designing complex surfaces that can be manufactured by AM. From a single surface feature design, arrays of said feature can be produced with variance in X, Y pitch and X, Y, Z dimensions. The variance can be normally distributed or defined by a mathematical expression. This allows for the creation of complex surfaces in a matter of minutes, whilst remaining user friendly. Arrays of one hundred pillars with varying diameters were fabricated by Two-Photon Polymerisation (2PP). The printed arrays were imaged with scanning electron microscopy and the diameters compared to the original CAD values to verify the ability of 2PP to accurately reproduce the designed variance. A case study is presented wherein surfaces are designed to increase biointegration of scaffolds for tissue engineering.

A case study is presented wherein surfaces are designed to increase biointegration of scaffolds for tissue engineering. Arrays of pillars and craters were produced by 2PP onto an existing part to determine the feasibility of 2PP as a method to improve existing scaffold surfaces.2. EXPERIMENTAL METHODOLOGY 2.1 Two-Photon Polymerisation2PP was chosen as the AM method to produce the designed surfaces as it is able to accurately build at the micron scale. A Nanoscribe™ Photonic Professional GT system was operated in Dip-In Laser Lithography (DILL) mode, a schematic of which is shown in Figure 2. The refractive index difference between the slide and photoresist is used to detect the slide surface and the build datum is set at this interface. To be found manually, practise builds were conducted at different build heights until the surface was found. Two-photon polymerisation occurs in the focal volume, the voxel, polymerising the photoresist.

Two-photon polymerisation occurs in the focal volume, the voxel, polymerising the photoresist. The piezoelectric scanning stage moves the slide relative to the fixed voxel.The surfaces were built from the photoresist OrmoComp®, an organic-inorganic hybrid polymer shown in other studies to be biocompatible [12]. A mean laser power of 30mW was used for structures outer shells, 20mW for internal hatching with a scan speed of 20mm/s. The OrmoComp® was post-cured by heat treatment at 140°C for 10 minutes. OrmoDev® followed by an application of 2-isopropranol was used to remove excess material.2.2 Verification TestsTest arrays were designed and built to determine if 2PP could accurately reproduce the design intent at the required scales. Three arrays of one hundred pillars were designed, with each pillar given a (cid:1850) domain value between 0.01 and 1. The diameter of the pillars in microns ((cid:1830)) was then determined by the functions listed in Table 1.

The diameter of the pillars in microns ((cid:1830)) was then determined by the functions listed in Table 1. The exponential functions were2425Piezoelectric 3D Scanning StageSlidePhotoresistVoxelBuild DatumObjective LensFigure 2: Schematic of the Dip-in Laser Lithography (DILL) mode used during two-photon polymerisation.chosen to be of the same form ((cid:1830) (cid:3404) (cid:1827)(cid:1857)(cid:3003)(cid:3025) (cid:3397) (cid:1829)) to determine if the 2PP process can accurately reproduce the coefficient values (cid:1827), (cid:1828) and (cid:1829) for each individual array. The arrays were built by 2PP in the method described in section 2.1 onto a glass slide. Scanning Electron Microscopy (SEM) was used to image the prints and from these ImageJ was used to determine the pillar diameters.

Scanning Electron Microscopy (SEM) was used to image the prints and from these ImageJ was used to determine the pillar diameters. The pillar diameters of the CAD design were then compared to those of the actual build.VT (1)VT (2)VT (3)Function(cid:1830) (cid:3404) 2(cid:1857)(cid:2870)(cid:3025)(cid:1830) (cid:3404) 2(cid:1857)(cid:2871)(cid:3025)(cid:1830) (cid:3404) 2(cid:1857)(cid:2870)(cid:3025) (cid:3397) 10Table 1: Functions used to define the pillar diameters ((cid:2160)) for the three verification test arrays.2.3 Surface Modification for Tissue Scaffolds2PP is limited by relatively slow build speeds. As such it may be more feasible for a tissue scaffold to be built by a quicker manufacturing method and the surface topography added in a second step by 2PP. This possibility was investigated by building arrays onto an existing part. 5x5mm Tri(Propylene Glycol) Diacrylate substrates, ~150µm in height, were used to simulate an existing tissue scaffold surface.

5x5mm Tri(Propylene Glycol) Diacrylate substrates, ~150µm in height, were used to simulate an existing tissue scaffold surface. The substrates were inkjetted by a PIXDROTM Toucan and the designed arrays were then printed onto the substrate surface by 2PP using the method described in section 2.1. Existing topographies that have been found to enhance osteointegration – pillars and craters – were used as the surface features [5]. Three arrays were designed, applying normal distributions to the X Y pitches, height and diameters of the features. Images of the CAD designs are shown in Figure 3. Each array was built multiple times, allowing for large areas of the substrate surface to be modified. SEM was used to image the prints and qualitatively evaluate the builds.2426TS (1)TS (2)TS (3)Figure 3: Designs of surface arrays to enhance biointegration of scaffolds. Features had their X Y pitches, diameter and heights normally distributed.

Features had their X Y pitches, diameter and heights normally distributed. The arrays were repeated in order to tile a large area of the substrate. TS (1): Pillar array on 112µm2 base. A sinusoidal pattern was superimposed to produce the concentric ringed pattern seen. TS (2): Crater array on 500x1137µm base. Mean X Y pitches are 128±14µm, diameter 54±5µm and height 9±1µm. TS (3): Pillar array on 175x195µm base. Mean X Y pitches are 10.0±0.2µm, diameter 1.00±0.05µm and height 2.0±0.2µm. Scale bars = 50µm.3. METHODOLOGY FOR SURFACE DESIGNThe developed methodology was based on the premise that it be able to apply and control variance to the surface features whilst being accessible for novice CAD users and not dramatically increasing design time. Grasshopper® was used to write the script for the surface design tool as it enabled synchronous modelling and the development of a basic Graphical User Interface (GUI). This allowed the stated objectives to be realised.

This allowed the stated objectives to be realised. In the script, the user can define the following parameters: X pitch, Y pitch and X Y Z bounding dimensions. These parameters can be varied for each surface feature either by a mathematical expression or by a normal distribution. The method for creating this quasi-normal distribution is visually demonstrated in Figure 4 .The user defines the mean (μ), and standard deviation (σ) for the chosen parameter, labelled (cid:1876) in this example. The script then generates N 2⁄ values for a standard normal distribution z, where	N is the total number of surface features in the array. The z values are shifted to (cid:1876) values by the following equation:(cid:1876) (cid:3404) 	μ (cid:3397) (cid:1878)(cid:2026)The values of (cid:1878) are generated in the following manner.

The following intervals were used: 0 to 2 in steps of 0.1 and 2 to 6 in steps of 1, giving 24 intervals in total.2427(1)(2)(a)(b)Figure 4: Diagrams illustrating how the script generates a quasi-normal distribution. (a) Equation (2) is used to determine the number of values within a given interval. Values of (cid:2162)(cid:4666)(cid:2208)(cid:4667) are linearly spaced in these intervals, seen pictorially as the vertical set of points. From these, values of (cid:2208) (horizontal set of points) are calculated by equation (3). Intervals are represented by the dashed lines. (b) A histogram of four hundred parameter values. The user defined mean is 10, and standard deviation is 2. The actual mean of the generated distribution is 10.00 and 1.97 respectively, to two decimal places.2428As this is the standard normal distribution the values of the integrals can be obtained from literature [13].

As the generated distribution is discrete, not continuous, there will be a discrepancy between the generated standard deviation and that defined by the user. By linearly spacing values of F(cid:4666)z(cid:4667), values of z are not linearly spaced within the equivalent interval, as seen in Figure 4 (a). At larger values of N, this results in the generated standard deviation more closely matching the users’ inputted value.The steps of the proposed methodology are shown in Figure 5. The user starts with an unmodified surface, assumed to be a flat horizontal plane. A surface feature, such as the crater shown in Figure 5 (i), is designed through any CAD package before being imported into the script as an STL file. Using slider inputs, the user can define the number of surface features in X and Y directions. The X and Y pitches can be defined and given variance, by either mathematical expression or a normal distribution.

The X and Y pitches can be defined and given variance, by either mathematical expression or a normal distribution. The resultant array of coordinates, seen in Figure 5 (ii), act as centre point for each surface feature. The X, Y and Z bounding dimensions of the features can be distributed in the same manner. The result is an array of surface features with defined variation, an example of which is shown in Figure 5 (iii).Then, if desired, an attractor point can be used to apply a user defined density increase to the array. An attractor point is placed on the array and a function is chosen using Grasshopper’s interactive ‘graph mapper’ input. The distance between the attractor and array coordinates are scaled between 0.01 and 1, with 0.01 being the closet coordinate to the attractor. The function is then applied to whichever parameter the user desires. An example of such a pattern is shown in Figure 5 (iv).

onto each array coordinateiii)Apply scaling function to desired S.F. dimensions (X,Y,Z)(iii)yesImprint pattern onto array?noDefine attractor parametersApply attractor pattern to desired S.F. dimensions (X,Y,Z)iv)Modified surface(iv)Figure 5: Flow Diagram of the proposed methodology to design complex surfaces. (i) Single surface feature design is imported into tool. (ii) Example of an array of coordinates that have had the X and Y pitches normally distributed. (iii) Example array of 400 craters with varying size and spacing. (iv) Example of a patterned array of 400 craters. The pattern was generated with an attractor point placed in the centre and sinusoidal function applied radially.24304. RESULTS 4.1 Verification ArraysFigure 6 shows the results of the verification tests. The printed arrays VT (1) (ii) and VT (3) (ii) have debris on some of the pillars. These pillars were not included in the analysis.

The printed arrays VT (1) (ii) and VT (3) (ii) have debris on some of the pillars. These pillars were not included in the analysis. As can be seen in the (iii) graphs, there is an offset between the CAD design and printed diameters. The print is always larger than the CAD design, though the fidelity is fairly uniform across length scales.VT (1)VT (2)VT (3)(i)(ii)(iii)Figure 6: Comparison of design and printed test arrays. Each column corresponds to the labelled test array. (i) Top views of design arrays. (ii) S.E.M. images of 2PP printed arrays. (iii) Comparison of measured diameters on the print and design arrays. The print data was fitted with functions of the same input form and the coefficients presented in Table 2. Scale bars = 50µm.2431DesignPrintFor VT (1) (iii), the small pillars approach 0µm in diameter.

Scale bars = 50µm.2431DesignPrintFor VT (1) (iii), the small pillars approach 0µm in diameter. The systems voxel size is ~200nm in x-y and ~1000nm in z, though this can be influenced by laser intensity, chosen focussing optics and photoresist material [14]. As such, diameters close the systems resolution limit cannot be accurately created. VT (2) (iii) shows as the diameter increases the offset between print and design is reduced. For VT (3) (iii) the offset appears consistent across all diameters, most likely due to the smallest diameters being much greater than the systems resolution. Whilst not quantitatively obtained, there will also be error introduced from the thresholding used during image analysis. This should be accounted for to more accurately determine any systematic error due to the Nanoscribe™ geometry interpretation algorithm.Table 2 shows the fit parameters used for the CAD and printed arrays.

Error on print parameters from 1σ confidence bound of the computed fit.4.2 Tissue Scaffold Surface ModificationsFigure 7 shows S.E.M images of the 2PP prints of the tissue scaffold surfaces. TS (1) print took ~7 minutes to build an area of 336µm2. TS (1) (iii) shows the concentric rings produced by the attractor and the variation in pillar height produced. Comparing to the CAD design in Figure 3, the top of the printed pillars became rounded, due to the resolution of the system and the ellipsoidal shape of the 2PP voxel.TS (2) build is 500µm x 1137µm	and took ~15 minutes to build. TS (2) (i) shows the designs base appearing to fade into the substrate. This is due to printing onto a non-flat substrate surface with a flat, fixed height build datum. This is shown schematically in Figure 8.

This is due to printing onto a non-flat substrate surface with a flat, fixed height build datum. This is shown schematically in Figure 8. This can lead to build errors as seen in TS (2) (iii), where misalignments between build sections have caused the crater halves to not join.2432(cid:1855)1011.50 (cid:3399) 0.06TS (1)TS (2)TS (3)(i)(ii)(iii)Figure 7: S.E.M images showing the printed surface arrays for enhanced biointegration. TS (1): (i) 3x3 grid of the repeated CAD design. (ii) Centre pattern. (iii) Side view of the centre right pattern. TS (2): (i) Top view of printed surface. The substrates surface can be seen to rise above the build datum where the base plate disappears. (iii) Error in the section alignment leads to parts of craters split over two build sections not joining up. TS (3): (i) 6x8 grid of the repeated CAD design. The design was split into four sections and as such consists of 12x16 squares with the design repeating every two squares.

TS (3): (i) 6x8 grid of the repeated CAD design. The design was split into four sections and as such consists of 12x16 squares with the design repeating every two squares. (ii) Section where substrate surface falls away from the build datum, leading to a split in the grid. (iii) Side view of pillars. Scale bars = 50µm.TS (3) took ~45 minutes to build the 1400µm x 1170µm design. As seen in TS (3) (i) sections have not joined properly on the left side of the array. At the edge of the substrate, the substrates surface drops dramatically below the build datum, leaving the printed sections here to drift freely down to the surface. This results in the large misalignments seen. Additionally a faded patch can be seen to the right of the build, due to the substrates surface rising above the build datum. TS (3) (ii) shows how similar faults can arise from building on top of existing surface features.

TS (3) (ii) shows how similar faults can arise from building on top of existing surface features. A closer view of the pillars is seen in TS 3 (iii).2433(a)(b)(c)Substrate SurfaceBuild DatumFeaturesFigure 8: Schematic showing how building on a non-flat substrate surface can lead to build errors. In (a) and (c), the substrate surface has risen above the build datum leading to features being built below, not on, the substrate surface, resulting in ‘faded patches’ as seen in Figure 7 TS (2) (i) and TS (3) (i). In (b) the surface has fallen below the build datum. Features are built unattached above the surface before drifting freely down to the surface. This can cause misalignments and deformations in the build, as seen in Figure 7 TS (2) (iii) and TS (3) (i) & (ii).Being close to the resolution of the system, the pillars have lost the defined ‘pillar shape’ of the CAD design. This is variation that is not designed for, and is a limitation of the manufacturing method.

This is variation that is not designed for, and is a limitation of the manufacturing method. These results emphasise the importance of the existing surfaces texture has when attempting to modify a surface with 2PP.5. DISCUSSIONThe proposed methodology simplifies what could otherwise be a complex process by using algorithmic programming to set up a large array of varied surface features. Grasshopper® working synchronously with Rhinoceros® makes the design process intuitive and user friendly to non-CAD experts. This methodology for surface modification can be applied to challenges other than those presented in this project, for example improving parts fluid dynamics or fatigue performance. Moreover, the methodology can be adapted for areas other than surface design, allowing for variance and mathematical expression to be easily incorporated into CAD parts.More work can be conducted in order to improve capabilities of the presented methodology.

Collision detection between surface features can be incorporated to ensure overlaps do not occur, which could lead to build errors during manufacture. As seen in the tissue scaffolds case study, error in the build can arise due to a non-flat substrate surface. Producing mathematical models or direct surface measurements of the substrate would allow the array to be mapped onto the surface. In addition the surface features are only squashed or stretched in X, Y and Z. By building the surface feature within Grasshopper® itself, it would allow for parts of the surface features to be scaled independently, instilling more detailed and controlled variation in the surface features.The verification tests and case study highlight the limitations and issues needed to be overcome using 2PP to manufacture complex surfaces. Surface features close to the resolution2434of the system are reproduced less accurately.

Surface features close to the resolution2434of the system are reproduced less accurately. Avoiding features of size ~1µm will improve the accuracy and reproducibility of the build. The systematic error between design and print should be quantified and included in the design process to allow for the features to come out at the desired size.Complications arose when attempting to modify the substrate surfaces; misalignment between build sections or missing features arose due to building on an undulating substrate surface with a single flat datum of fixed height. The Nanoscribe™ Photonic Professional GT has an interface finder capability, which did not work in this case study due to an insufficient difference in the refractive index of TPGDA and OrmoComp®. Appropriate choices of material could allow this function to be implemented, allowing each build sections datum to be located onto the substrates surface.6.

Appropriate choices of material could allow this function to be implemented, allowing each build sections datum to be located onto the substrates surface.6. CONCLUSIONA methodology for conducting complex surface design has been presented. A script was written that used algorithmic programming to incorporate variation into a surface, allowing for arrays of hundreds of uniquely sized and spaced features to be created from a single surface feature design. This can be built by AM, allowing complex surfaces to be implemented without an unreasonable increase in design or lead time. The ability to create complex, functional surfaces has uses in many design areas, for example improving parts fluid flow or fatigue performance. This methodology enables novel surface designs to be feasibly created for these applications.

This methodology enables novel surface designs to be feasibly created for these applications. Further work can be undertaken to add more capabilities to the script, such as the ability to scale parts of surface features independently, creating a more powerful design tool.Verification tests demonstrate that 2PP is a viable process to precisely build the surface design at the microscale. Whilst close in accuracy to the original CAD sizes, this could be improved by quantifying any systematic error and then offsetting this within the actual CAD design. A case study for the surface modification of tissue scaffolds to enhance biointegration was conducted. Mesoscale arrays were printed onto an existing substrate, demonstrating the feasibility of such modification. Errors with array alignment due to a non-flat substrate surface need to be addressed to improve accuracy and reproducibility.

Errors with array alignment due to a non-flat substrate surface need to be addressed to improve accuracy and reproducibility. With future developments, the presented methodology may enable healthcare professionals to selectively design custom scaffold surfaces prior to manufacturing, allowing for higher performing tissue scaffolds without huge rises in cost and lead time.Whilst in the preliminary stages, the methodology demonstrates the freedom of design and complexity achievable with AM surfaces with the appropriate CAD tools. Such algorithmic methods can open up new design possibilities, having significant future applications for surface modification to improve part performance.24357. ACKNOWLEDGEMENTSThanks to the EPRSC (Engineering and Physical Sciences Research Council) Centre of Doctoral Training in Additive Manufacturing and 3D Printing for support and funding.

mposite Modeling and Analysis of FDM Prototypes for Design and Fabrication of Functionally Graded Parts Longmei Li, Q. Sun, C. Bellehumeur1, and P. Gu 2 Department of Mechanical and Manufacturing Engineering 1Department of Chemical and Petroleum Engineering The University of Calgary, Calgary, Alberta, T2N 1N4, CANADAAbstractSolid Freeform Fabrication technologies have potential to manufacture parts with locally con- trolled properties (LCP), which would allow concurrent design of part’s geometry and desired properties. To a certain extent, Fused Deposition Modelling (FDM) has the ability to fabricate parts with LCP by changing deposition density and deposition orientation. To fully exploit this potential, this paper reports a study on the mechanical properties of FDM prototypes, and related materials and fabrication process issues. Both theoretical and experimental analyses of mechanical properties of FDM parts were carried out.

Both theoretical and experimental analyses of mechanical properties of FDM parts were carried out. To establish the constitutive models, a set of equations is proposed to determine the elastic constants of FDM prototypes. An example is provided to illustrate the model with LCP using FDM. Key words: Fused Deposition Modeling, Composite Mechanics, Locally Controlled Properties.1. IntroductionWith recent advances in Solid Freeform Fabrication (SFF) technologies, it is possible to use these technologies to produce parts with desired mechanical properties. The ability to fabricate parts with locally controlled properties (LCP) would create opportunities for manufacturing of a whole new class of parts with graded compositions [Jackson et al., 99] and [Cho et al., 01]. Fused Deposition Modeling(FDM) has the potential to produce parts with LCP by changing deposition density and deposition orientation. A few examples with LCP are shown in Figure 1 b), c), d), e).

Fused Deposition Modeling(FDM) has the potential to produce parts with LCP by changing deposition density and deposition orientation. A few examples with LCP are shown in Figure 1 b), c), d), e). FDM processes fabricate prototypes by extruding a semi-molten ﬁlament through a heated nozzle in a prescribed pattern onto a platform layer by layer, as in shown Figure 1 a). After the semi-molten material is deposited onto the worktable it begins to cool and bond to the neighboring material diffusely. The bonding between the individual roads of the same layer and of neighboring layers is driven by the thermal energy of semi-molten material.

The bonding between the individual roads of the same layer and of neighboring layers is driven by the thermal energy of semi-molten material. FDM processes use several types of materials, such as Acrylonitrile-butadiene-styrene (ABS) and wax, to build conceptual and/or functional prototypes.Most of the existing research work in FDM is primarily directed towards the development of new materials or techniques for material deposition [Agarwala et al., 96] [Gray IV et al.,98], [Yardimci et al., 96]. The mechanical properties of ABS ﬁlament and FDM parts were also investi- gated by researchers such as [Comb et al., 94], [Kulkarni and Dutta, 99] and [Rodriguez Matas, 99]. ABS parts made by FDM processes essentially consist of partially bonded ABS ﬁlaments. Mechanical properties of FDM prototypes are governed by their meso-structures that are determined by manufacturing parameters.

Mechanical properties of FDM prototypes are governed by their meso-structures that are determined by manufacturing parameters. By selecting manufacturing parameters, such as deposition directions in layers and gap sizes between ﬁlaments, FDM processes can achieve LCP and produce prototypes with desired properties. To fully exploit this potential, issues concerning the mechanical properties of FDM parts and the fabrication process parameters, should be investigated. This paper describes a study on the mechanical properties of FDM prototypes.2To whom all correspondences should be addressed: gu@enme.ucalgary.ca187Figure 1: FDM process and the parts fabricated with graded mechanical propertiesThe remaining sections of this paper consist of the following parts: Section 2 is devoted to the modeling of the mechanical properties for design of FDM parts; experimental validation is given in Section 3.

Section 4 provides an example of the prototypes with graded mechanical properties using LCP. The ﬁnal section summarizes the main conclusions.2. Modeling and Analysis of Mechanical Properties of FDM PrototypesThe FDM prototypes are orthotropic composites of ABS ﬁlaments, bonding between ﬁlaments and voids. They can be viewed and analyzed at different levels and on different scales. At macro level, they are studied as laminates of bonded laminars. At micro level, the properties of each laminar are functions of properties of ﬁlaments and void density.2.1 Lamination Analysis of FDM Parts. To concurrently design mechanical geometries and properties, it is important to determine the effective stiffness so that the model can be used to predict the mechanical behaviors of the parts. Filament orientation (cid:18) is deﬁned as the angle be- tween the loading direction and longitudinal direction of ﬁlaments of each lamina.

General orthotropic lamina refers to a lamina with ﬁber orientation other than o. Although, o or the stiffness matrix for general orthotropic lamina is fully populated with 6 elements. These six elements are functions of the four independent elements in matrixare the in-plane stress and strain respectively.and]Q[(cid:15)]](cid:27)[[ijij22, major Poisson’s ratio (cid:23)12, and shear Modulus G[Q]090[Mallick, 93].A laminate is constructed by stacking a number of laminas in the thickness direction. The188(1)overall properties can be modeled by the Lamination Theory [Mallick, 93]. As such, in macrome- chanical approaches, the stiffness matrix of a laminate is composed by properties of every lamina according to the Lamination Theory, while the in-plane elastic behavior of a unidirectional lamina may be fully described in terms of four elastic lamina properties including E 12.

For the design purpose, it is desirable to have reliable predictions of lamina properties as a function of constituent properties and geometric characteristics. One of the objectives of micromechanics is to obtain such relationships.11, E22, G12 and (cid:23)2.2 Micromechanics of Material with Voids. A variety of methods have been used to pre- dict properties of composite materials as functions of constituent volume ratio and geometric packing parameters. Tsai [Tsai, 64] conducted an analysis based on variational principles in which contiguity of the ﬁbers and ﬁber misalignment were considered. The ﬁlament contiguity was incorporated in theoretical prediction E 11.

The ﬁlament contiguity was incorporated in theoretical prediction E 11. The set of the equations, when being reduced to material with voids, is depicted below.22, and the ﬁber misalignment was considered in EThe longitudinal Young’s modulus can be described by the equation:(2)where E is the Young’s modulus of the polymer, (cid:26) (normal to the ﬁbers), and k is an empirical factor that takes into account the fact that the ﬁlament may not be exactly parallel or perfectly straight. The prediction of transverse modulus, E1 is the void density in Plane122, takes form:(3)where (cid:23) is the Poisson’s ratio of the polymer. The effective in-plane shear modulus, G by:12, is given(4)where G is shear moduli of the polymer. And the in-plane Poisson’s ratio, (cid:23) following relation:12, is given by the(5)Only the void density in the plane perpendicular to the ﬁbers is required to compute the effective properties.

The value of ﬁber contiguity factor C can be assigned to two extreme cases, (1) all ﬁlaments are isolated, C . The effective Poisson’s ratio, 12, is found to be the same as that of polymer when C . It is the same as Hashin and Rosen’s conclusion [Hashin and Rosen, 64].; (2) all ﬁbers are contiguous, C= 0= 1(cid:23)= 0Tsai’s equations may be suitable for materials with cavities, but they could be used for FDM parts only under the assumption of perfect bonding among ﬁlaments. However, this assumption is not valid for the FDM process according to the ABS sintering experiments that were conducted in this research.189Figure 2: Fiber volume fraction in the transverse direction2.3 Modeling of Elastic Constants for FDM Prototypes. The previous section demonstrated that FDM prototypes are composites of partially bonded ABS ﬁlaments and voids.

The previous section demonstrated that FDM prototypes are composites of partially bonded ABS ﬁlaments and voids. Because of imperfect bonding, A new set of equations is proposed to calculate the elastic constants.When loaded in longitudinal direction, a unidirectional FDM part is an aggregate of ABS ﬁlaments. Longitudinal Young’s Modulus can be predicted well by the rule of mixtures. Therefore, Longitudinal Young’s Modulus is:(6)where E is Young’s Modulus of P400 ﬁlament, and (cid:26) 1 is the area void density in the plane normal to ﬁlaments. In the case of transverse normal loading, the bonding among the ﬁlaments is the load carrier. The following equation is proposed for Transverse Modulus:(7)where (cid:16) is an empirical factor that takes into account the bonding strength. (cid:26) along the transverse direction.

The following equation is proposed for Transverse Modulus:(7)where (cid:16) is an empirical factor that takes into account the bonding strength. (cid:26) along the transverse direction. Refer to Figure 2, the deﬁnition of (cid:26)2 is:2 is linear void fraction(8)Term ( in-plane shear modulus takes form:2) is the ratio of bonding length to overall dimensions. Similarly, the prediction of the(cid:0)(cid:26)1(9)where G is the Shear Modulus of P400 ﬁlament, and the in-plane Poisson’s ratio, (cid:23) that of ABS ﬁlaments, (cid:23):12, is same as(10)Comparing Equations 6 through 10 with Equations 2 to 5, the signiﬁcant differences are in the predictions of E 2 is introduced to count the potential load carrying dimension, whereas (cid:16) results in the ratio of the effective load carrying media.

Since the estimation of these elastic constants requires the void density in the plane perpendicular to ﬁlaments 2, the calculations of the void densities are the topic of the next section. 1 and bond length ratio (cid:26)22 and G12. Void density (cid:26)(cid:0)1(1(cid:26))2.4 Micro-scopic Images Analysis. Theoretical calculations of strength, modulus, and other properties for a FDM part are based on the the void densities (cid:26) 2. Therefore the cross- sections of FDM specimens are studied under a microscope. Unidirectional cubic specimens of dimension mm were made with different gap sizes. They were cut in the middle plane1 and (cid:26)190Table 1: Results of digitized image analysisGap Size( -50.0 -25.4 0 100)(cid:26)10.0343 0.0434 0.0799 0.1990(cid:26)20.2776 0.3419 0.4038 1Figure 3: Meso-structures of different gap settings. a) gmm; b) gmm; d) gmm and e) gwhich is normal to ﬁlaments.

a) gmm; b) gmm; d) gmm and e) gwhich is normal to ﬁlaments. Then the photos of meso-structures were taken under a microscope, as shown in Figure 3. They were analyzed by using image processing software. Void densities 2 were calculated, and the results are listed in Table 1. From the observation and the 1 and (cid:26) image analysis, it can be concluded that the voids appear periodically when the negative gap is small. When the negative gap reaches (cid:0) mm, the dimensional accuracy and surface quality of the specimen become very poor, and the periodicity can rarely be observed.(cid:26)0:13. Experimental ValidationThe theoretical stiffness model developed previously is a function of deposition densities and deposition orientations. To validate the model and method, the experimental analysis was carried out. The experiments tested the orthogonal-ply specimens, and the experimental results were compared with those of constitutive modeling described in Section 2.

The experiments tested the orthogonal-ply specimens, and the experimental results were compared with those of constitutive modeling described in Section 2. All specimens were manufactured by the FDM2000 machine.

All specimens were manufactured by the FDM2000 machine. The elastic constants of P400 ﬁlaments were cited from work of Rodriguez Matas [Rodriguez Matas, 99], listed in Table 2, as the inputs of the theoretical calculations.Table 2: Elastic Properties of P400 ﬁlamentElastic Constant Young’s Modulus, Shear Modulus,Value(cid:6)15M P a191mm; c)1800170016002m m N/1500,x x E14001300120005101520 25 30 Raster angle q35404550Figure 4: Young’s modulus vs deposition Orientation in orthogonal-ply specimensTable 3: Orthogonal-ply specimen test results Difference Percent 3.1 2.4 5.0 7.1SpecimenYoung’s ModulusM P aExperimental Theoretical[0/90] [15/-75] [30/-60] [45/-45]1772.9 1623.4 1440.4 1369.61718.7 1584.7 1368.2 1272.7The stiffness matrix of orthogonal-ply specimens can be constructed according to the procedure described in Section 2.

As a result of the laminate modeling, the in-plane Young’s modulus of the orthogonal-ply symmetric laminates with different [(cid:18)/ ] ﬁber orientations is a non-linear function of the deposition angle, as shown in Figure 4. It can be observed from the ﬁgure that the laminate with raster angles of approximately [45/-45] yields the minimum modulus of elasticity. The laminate with raster angles of [0/90] has the highest Young’s Modulus. The experiments with the orthogonal-ply symmetric specimens were tested to determine the same in-plane moduli. The specimens with raster angles of [0/90], [30/-60], [45/-45], and [15/-75] were manufactured with the same thickness of 11 layers. These experimentally determined values were then compared with the theoretically predicted values to evaluate the effectiveness of the proposed laminate modeling approach, in Table 3 and marked as "(cid:3)" in Figure 4.

The experiment data were in good agreement with the results of laminate modeling.o(cid:0)(cid:18)(90)4. ExampleThe purpose of the study of mechanical behaviors of FDM prototypes is to establish the models for design and fabrication of the parts with the desired mechanical properties through local composition control. This can be carried out with FDM process by changing deposition density and orientations to obtain the requried properties.The design of the example part with graded mechanical properties is shown in Figure 5a). Assume that the geometrical and functional constraints require the smaller section with equal or higher stiffness than that of the larger section areas. Therefore, the properties are gradient along192Table 4: Conﬁguration of the axisLength()Deposition strategyoGap ()20 [0/0] -25.4 2137.720 [0/90] -25.4 1718.720 [45/-45] -25.4 1272.7Figure 5: An axis with gradient property.Axis X.

Different deposition strategies and deposition densities were used for different sections of the part to meet the requirements as listed in Table 4. Young’s modulus E of each section calculated according to the stiffness modeling are listed in the table. If an external axial force P is applied at each end, the overall elongationxxL is calculated using the following equation:where L is the cross-sectional area of the section. With locally controlled deposition density and deposition orientation, it is possible to make the section with smaller cross-sectional area as stiff as, or even stiffer than, the section with bigger cross-sectional area.is the length of each section, Eis the Young’s Modulus of the sectoin, and Aiixx5. Concluding RemarksThe design and fabrication of products with LCP would make it possible to create a new class of products with desired mechanical and material characteristics.

Concluding RemarksThe design and fabrication of products with LCP would make it possible to create a new class of products with desired mechanical and material characteristics. This paper reported a study on modeling of in-plane mechanical properties of FDM prototypes relating to fabrication parameters. Based on both the theoretical and the experimental analyses, a set of new equations of calculating the elastic constants were proposed and validated, which were used to determine the constitutive models of FDM parts. An example was presented to illustrate the design and fabrication of the FDM prototypes with required mechanical properties by locally controlled deposition density and orientations. Different deposition densities, angles and their combinations can be employed in producing the required stiffness properties of manufactured parts.

Different deposition densities, angles and their combinations can be employed in producing the required stiffness properties of manufactured parts. The further study is currently underway in investigating void geometry and density with more manufacturing parameters.193(11)x)AcknowledgementsThe ﬁnancial support for this research was provided by the National Sciences and Engineering Research Council of Canada (NSERC) through Research Grant OGP0192173. The authors would like to thank Brandon Furguson and Art Moehrle for their assistance with experiments performed in the research.References[Agarwala et al., 96] M. K.Agarwala,A.Bandyopadhyay,R.vanWeeren,N. A.Langrana,A.Safari,andS. C. Danforth, Fused Deposition of Cerameics (FDC) for Structural Silicon Nitride Components, Solid Freeform Fabrication Symposium, Austin, TX, Aug. 8th, 1996.[Gray IV et al.,98] Robert W. Gray IV, Donald G.

MPOSITES WITH GRADIENT PROPERTIES FROM SOLID FREEFORM FABRICATIONV.R. Gervasi, R.S. Crockett Rapid Prototyping Center Milwaukee School of Engineering, Milwaukee, WIAbstractTetraCast is a build style developed by Milwaukee School of Engineering involvingstereolithography patterns produced with an open cellular structure inside a surface shell. Composites are created using this pattern as a host fora filler material, generally epoxy matrices loaded with various fibers or microspheres. Regions within a singleobjectrrray be separated by thin barriers, allowing filling with different matrix materials. to create regiollsof differing. local properties.

Regions within a singleobjectrrray be separated by thin barriers, allowing filling with different matrix materials. to create regiollsof differing. local properties. The internal structure can also be continually graded in thickness to prodUl::e composites with properties ranging from that of the filler material to that of the TetraCast material (currently stereolithography epoxy or FDM ABS).IntroductionFunctionally gradient materials (FGMs) transform from one material (part A) to another(part B), unlike traditionally bonded materials that have a distinctive seam. The goal ofFGMs is generally to produce materials able to withstand extremely harsh environments. Examples include surviving temperatures as high as 21 OOiK and temperature gradients, within the FGM, as great as 1600 K over a very short region. FGMs also need to handle these temperatures while maintaining mechanical functionality. To accomplish this, the.

FGMs also need to handle these temperatures while maintaining mechanical functionality. To accomplish this, the. benefits of two or more materials are combined in aFGM to produce a non-traditional engineeringmaterial1 • A good example is a FGM consisting of metal and ceramic, as illustrated in figure 1. The properties of the metal componentare high streIlgth,ductility,conductivity (electrically&thermally) and low chemical resistance..The. properties of the ceramic arehighhardnyss, insulating, brittle, chemicallyFigure 1. Metal-CeramicFGM729resistant, and abrasion resistant. Now, if the advantages of both are combined in a FGM, the result is a material with high mechanical strength, conductivity, corrosion resistance, and abrasion resistance.

Now, if the advantages of both are combined in a FGM, the result is a material with high mechanical strength, conductivity, corrosion resistance, and abrasion resistance. Combinations being tested include: metal-ceramic, ceramic-ceramic, carbon-carbon, carbon-ceramic, and polymer-polymer.Industries with applications for FGMs include aerospace, manufacturing, mining, nuclearpower, architecture, automotive, biomedical, and electronics. In aerospace, leading edges and combuster cans could benefit from FGM properties. Whereas in manufacturing and mining, abrasive tools could be produced with a combination of wear resistant surfaces and shock absorbing surfaces such as those found in FGMs. Biomedically, bone implants could be generated with bio-compatible materials on one surface and mechanically functional materials on another surface.The Production of FGMs is quite challenging, and numerous techniques are being testedand used.

A technique, such as vapor deposition, creates gradient materials at the atomic or molecular level. Other techniques, on the micro-scale, produce gradients by combining powders or laminates to produce FGMs. A new approach uses a meso-scale gradient morphology. This morphology, discussed in this paper, is called Gradient Tetracast.BackgroundBefore discussing the approach to this research, it is helpful to give some background on the TetraCast Build Style, the Gradient Tetracast Morphology, and the Rapid Composite Process.TetraCase is a SFF build style used to produce lattice-filled patterns for the Rapid Composite Process and investment casting. TetraCast patterns have many of the same benefits asQuickCast™ patterns. TetraCast uses an internal three-dimensional tetrahedron morphology modeled after the molecular bond geometry found in diamonds. The TetraCast internal lattice offers the outer surface or skin sufficient support with minimal density.

The TetraCast internal lattice offers the outer surface or skin sufficient support with minimal density. Figure 2 illustrates the TetraCast Lattice.Figure 2. TetraCast LatticeGradient Tetracast (GTC) is a three-dimensional morphology, transforming from onematerial to another, in a controlled manner. This morphology can be likened to M.S. Esher art work in three dimensions. To create Gradient Tetracast, spacing between standard TetraCast units can remain constant while the branch thickness is increased in a controlled manner (as shown in figure 3). This controlled increase in branch thickness results in a proportional change in volume percentage of material combinations. In theory, this volume percentage will transform730from 100 percent of material A to 100 percent of material B at any rate, in any direction. With Gradient Tetracast the transformation is three dimensional and mechanically interlinked. GTC becomes two TetraCast lattices with varying branch thickness.

With Gradient Tetracast the transformation is three dimensional and mechanically interlinked. GTC becomes two TetraCast lattices with varying branch thickness. Part A is the mirrored negative of part Band GTC can be stacked as shown in figure 4.Front ViewFigure 3. Gradient TetracastFigure 4. Stacked Gradient TetracastThe Rapid composite Process3 (RCP) is used to create composite parts from TetraCastpatterns. With this method, the TetraCast pattern becomes part ofthe final composite, as it both731definesthe net-shape ofthe desire.d geometry and embodies the composite material during matrix solidification. The composite material is introduced into the TetraCast pattern as chopped fibers, flakes, and/or particles suspended in a liquid matrix (typically epoxy).

The composite material is introduced into the TetraCast pattern as chopped fibers, flakes, and/or particles suspended in a liquid matrix (typically epoxy). After the TetraCast pattern is completely filled, the matrix solidifies,resulting ina net-shape functional composite part.Throughthislllethod,thecol11posite material properties .can be tailored to meetfirstexperil11ent.TheRapid C0111positeprocess functions well to fill small features such as those foundinthe first GTCsal11ples.ObjectiveThe main obj~ctive of this research was to introdu~ea new gradient morphology that usesGTC as the transitionaI·structure.•• To. accomplish this· a simpleGTCmaterial was produced consisting.ofpart A (SLA photopolymer)transitioning into part B (glass filled epoxy). Sample Functionally Gradient Materials were produced within a class III tensile geometry profile.ApproachThe approach to studying GTC began with the production ofplastic-compositeGTCtensile samples. glass filled epoxy.

glass filled epoxy. The main objective was to produce a visual sample of GTC to be used for discussion.andanal~sis.Tfie moduillsofelasticity as a function of percentage composition was also· analyzed. modeling, SLAproduction, RCP, sample slicing, and Testing.'The material combination was 50 percent SLA photopolymer and 50 percent. Thea~proachus~d toprodtlcethe first GTC sample consisted of: CADThe GTCtensile bar CA.D modelwas>designedto include five distinctive TetraCastpercentage combinations to form a gradient. .. The first TetraCast region was less than 10 percent by volume increasingto greaterthan 90 pergentinthe finalregio~. Each TetraCast section was 0.45 inches in thicknessto allow machine stockfor later slicing and sanding. All five TetraCast regions were combined within atfiin shell. One hundred. percent SLA photopolymer and composite samples were produced. separately.Two GTC samples were produced oU an SLA 250, usingACESbuildstyle.

One hundred. percent SLA photopolymer and composite samples were produced. separately.Two GTC samples were produced oU an SLA 250, usingACESbuildstyle. Uncuredphotopolymer was removed in a centrifuge, leaving a 50 percent graded air void and 50 percent graded photopolymer. SLA Photopolymer was chosen because ofthe fine detail, ease ofuncured resin removal, clarity, and homogeneous mechanical properties (specifically Young's Modulus of Elasticity).TheRCP was used to fill the 50 percenfgraded void with glass filled epoxy as shown inFigures 5. Glass filled Epoxy was chosen for its high modulus and grey color.Samples were sliced, polished, and tested for Young's modulus of elasticity. Figure 6illustrates the sectioned GTC sample.ResultsIDiscussionAs shown in Figtlre5, agradient material was suc;ce~)sfll1ly produced using the morphology.

Figure 6illustrates the sectioned GTC sample.ResultsIDiscussionAs shown in Figtlre5, agradient material was suc;ce~)sfll1ly produced using the morphology. Thetensilesarnples produced withGTC had a volllfrle percentage rangefrom 95 percent photopolyrner/5percent composite to 28 percent photopolymer172 percent composite. Low volume percentage photopolymer·sections collapsed during Rep filling and were not used. This section. would have been 95 perc~nt composite/ 5 percent Photopolymer..733After sample slicing (figure 6), tensile testing was performed to determine the modulus of each section. Figure 7 illustrates the resulting stress-strain curves, clearly displaying an increase in stiffness as percentage volume of composite increased. Figure 8 illustrates the change in modulus as a function of material composition. The samples were expected to act similar to that of a particulate composite. The Upper Limit shown is the predicted moduli applying the rule of mixtures.

Gradient TetraCast has the potential to be stackable, allowing for two or more materials in Functional Gradient Materials. Gradient TetraCast is a continuous structure providing thermal and electricaLconductivitythrough metallic portions of Functional Gradient Materials.Future WorkMuch work is required before the usefulness ofGradient TetraCast for FunctionalGradient Materials is determined. In the future, metal-ceramic Functional Gradient Materials will be attempted along with other suitable material combinations. Applying Finite Element Analysis (FEA) to both the mechanical analysis and Gradient TetraCast unit design will also be considered for optimizing design and production.

Applying Finite Element Analysis (FEA) to both the mechanical analysis and Gradient TetraCast unit design will also be considered for optimizing design and production. The production of simple to complex shapes will be attempted.AchnowledgementsThe authors gratefully acknowledge the support of the Rapid Prototyping Consortium onthis and other research.The Authors also extend their thanks to the Rapid Prototyping Center Staff at MilwaukeeSchool of Engineering for their contribution to thisresearch.Special thanks to Tom Bray, Dean ofApplied Research, for his SFF insight andencouragement towards.gradient material synthesis.Bibliography 1. Holt, J. Birch, Mitsue Koizuma, Toshio Hirai, Zuhair A. Munir, "Functionally Gradient Materials," American Ceramics Society, 1993, chaps 1 & 2 2. Gervasi, Vito R., Dr. Daniel A. Brandt, StevenD. Shaffer, King Lim, TetraCast SLA Build Style, International Conference on RP, 1997, pp 309-317 3.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCompressive and Bending Performance of Selectively Laser Melted AlSi10Mg StructuresD. Murphy, O. Fashanu, M. Spratt J. Newkirk and K. Chandrashekhara Missouri University of Science and Technology, Rolla, MO 65409H. Misak and D. Klenosky Spirit AeroSystems Wichita, KS 67278AbstractSelective laser melting (SLM) is a widely used additive manufacturing technique that effectively manufactures complex geometries such as cellular structures. However, challenges such as anisotropy and mechanical property variation are commonly found due to process parameters. In a bid to utilize this method for the commercial production of cellular structures, it is important to understand the behavior of a material under different loading conditions.

In a bid to utilize this method for the commercial production of cellular structures, it is important to understand the behavior of a material under different loading conditions. In this work, the behavior of additively manufactured AlSi10Mg under compression, bending, and tension loads was investigated. Vertical and horizontal build directions are compared for each type of loading. Specimens were manufactured using the reduced build volume (RBV) chamber of the Renishaw AM 250 SLM machine.1. INTRODUCTIONAdditive manufacturing as a whole began as a process to build small-scale prototypes, but over the last decade has developed into a widely accepted and utilized method of manufacturing. Selective laser melting (SLM) is one of the many types of metal additive manufacturing that has been well developed in this time frame. Some benefits of SLM manufacturing include reduced cost, design flexibility, and improved mechanical property performance.

Some benefits of SLM manufacturing include reduced cost, design flexibility, and improved mechanical property performance. However, large scale implementation of this process has yet to develop because of problems such as anisotropic behavior and inconsistent mechanical properties. Process parameters consist of material selection, laser parameters, and build parameters. All of these parameters have been shown in previous studies to affect microstructure, density, and mechanical properties of the SLM part. Understanding the effects of parameter changes is an important research topic to allow for the continued success of the SLM manufacturing technique [1-8].Aboulkhair et al [1] focused their study on reducing the porosity of parts by optimizing the laser parameters. The study claims to have produced dense parts with relative densities of 99.8%.

The study claims to have produced dense parts with relative densities of 99.8%. Brandl et al [2] varied the process parameters of build orientation, build plate temperature, and post process heat treatment to investigate their effects on high cycle fatigue, microstructure, and fracture type. It was found that the post process heat treatment had the most significant effect on each of these outputs. In another study done by Aboulkhair et al [3], the focus was to determine the effect on microstructure, tensile strength, and microhardness after a T6 heat treatment. While some mechanical properties increased (microhardness), others decreased (tensile strength). The SLM printed AlSi10Mg out performed a comparable cast aluminum alloy. Biff et al [4] researched2177the effect process parameters had on the behavior of the Si inside the fully dense AlSi10Mg part.

Biff et al [4] researched2177the effect process parameters had on the behavior of the Si inside the fully dense AlSi10Mg part. It was found that the energy density of the laser was the most impactful process parameter to the development of Mg2Si precipitation. Other mechanical properties like creep have been investigated by Read et al [5] after a laser parameter optimization was conducted. It again was found that the SLM manufactured aluminum parts out-performed the conventionally manufactured cast alloys of similar composition. A study looking at the anisotropic behavior of SLM printed 304L stainless steel due to build parameters was done by Fashanu et al [6].As mentioned, one of the biggest advantages to using the SLM process is the ability to fabricate complex shapes and geometries. Of these complex shapes, periodic lattice structures are of major interest for their ability to increase the strength to weight ratio of a part and that is the future goal of this current work.

Of these complex shapes, periodic lattice structures are of major interest for their ability to increase the strength to weight ratio of a part and that is the future goal of this current work. Maskery et al [7] fabricated and tested AlSi10Mg truss structure samples with uniform and graded density for their mechanical properties. Samples were also tested for heat treatment effects and found that heat treating increases the strength of the samples. Dong et al [8] studied lattice structures as well but wanted to see how the build orientation would affect the geometric accuracy, microstructure, and mechanical properties of the as built complex geometric parts. It was determined the geometric accuracy decreased as the strut angle went from increased from 35.5º to 90º (vertically built). On the other hand, the tensile properties were seen to increase as the angle increased since the parts had less porosity when built vertically compared to being built at an angle.

On the other hand, the tensile properties were seen to increase as the angle increased since the parts had less porosity when built vertically compared to being built at an angle. In the current study, the mechanical properties of AlSi10Mg coupons manufactured with the SLM process were investigated. Solid coupons with varied build orientation were tested under compression, bending, and tension loading to give an idea of what behaviors are to be expected from geometrically complex parts. The coupons’ yield strengths were evaluated from the developed stress-strain curves and compared with previous studies.2. EXPERIMENTAL METHODOLOGYThe experimental portion of this study consists of three experiments: first, solid sample compression tests with varied build orientation; second, solid sample 3-point bend testing with varied build orientation; third, mini-tensile testing was conducted for comparison to previous studies.

The 0.2% offset yield strength is reported for each test, as well as the nonlinear stress- strain curves.2.1 Fabrication of AlSi10Mg specimens with the SLM processFor the three experiments, the parts were fabricated using the RBV chamber in the Renishaw AM 250 machine at Missouri University of Science and Technology. The following describes the powder used, the SLM fabrication process, and the density measurements of the fabricated parts:1. Powder characterization: AlSi10Mg is the powder used for the SLM process in this study. The size distribution of the powder particles was determined and summarized in Figure 1. The size distribution was determined by automatically counting over 8000 particles using an ASPEX scanning electron microscope (SEM) system located at Missouri S&T.

Pre-processing: Test coupons were modeled in three dimensions using CAD software with a geometry correlating to the testing method. Models are exported as a Stereo Lithography (STL) file to Magics. The Magics software creates the laser toolpath for the Renishaw machine and specifies build parameters of the SLM parts. The STL file is then provided to the SLM machine for fabrication of the three types of test specimen.3. Fabrication: The Renishaw AM 250 machine was used to manufacture the SLM parts, using the RBV chamber in an Argon atmosphere. The powder is melted via a laser heat source in a layer by layer process. The laser parameters used during the part fabrication are shown in Table 1.

The powder is melted via a laser heat source in a layer by layer process. The laser parameters used during the part fabrication are shown in Table 1. Table 1: SLM laser parameters used for all samples [3]Laser PowerLayer ThicknessScan SpeedHatch SpacingScan StrategyPoint DistanceExposure Time200W(cid:21)(cid:24)(cid:3)(cid:541)(cid:80)570 mm/s(cid:20)(cid:22)(cid:19)(cid:3)(cid:541)(cid:80)Chessboard(cid:27)(cid:19)(cid:3)(cid:3)(cid:541)(cid:80)(cid:3)(cid:20)(cid:23)(cid:19)(cid:3)(cid:541)(cid:86)21794. Post-processing: After fabrication, the test coupons were separated from the build plate by electric discharge machining (EDM). The parts were then machined to proper testing dimensions (listed in their test specification section 2.2-2.4). The compression samples were machined on a lathe. The bending samples were machined on a shaper, and the mini-tensile samples were machined again using the EDM.5.

The compression samples were machined on a lathe. The bending samples were machined on a shaper, and the mini-tensile samples were machined again using the EDM.5. Density measurements: After machining was completed, the compression and bending samples’ densities were measured using calipers and a scale. The blocks, that the mini-tensile samples were cut from, were measured for their density before machining by Archimedes method. The average densities are recorded in Table 2. The densities of the samples were considered acceptable for testing to proceed. Table 2: Density measurements for all tested samplesSample TypeBulk Density (g/cm3)Compression2.69 ± 0.02Bending2.64 ± 0.02Mini-Tensile Blocks2.59 ± 0.01Theoretical2.672.2 Specifications and procedure for compression testingThe compression samples were fabricated first as octagons on the SLM build plate to allow for ease of printing and post processing.

After removal from the build plate using the EDM they were machined with a lathe into cylindrical samples to align with ASTM E9-19 Standard Test Methods of Compression Testing of Metallic Materials at Room Temperature [9]. The dimensions of the build plate octagonal prisms and machined cylinders are shown in Figure 3.15mm6.5mm30mm13mmFigure 3: Compression sample built and machined dimensionsThe length to diameter ratio is shown to be about 2:1. The build orientation was varied for the compression samples. Figure 4 shows the three different build orientations (x, y, and z) used2180while fabricating the compression samples as well as images of the machined samples.

Figure 4 shows the three different build orientations (x, y, and z) used2180while fabricating the compression samples as well as images of the machined samples. There were three samples made for each orientation, but as can be noted by the red circle in Figure 4 an issue with the final z build direction sample was seen due to complications using the RBV.Figure 4: a) Compression sample build plate with labeled build orientations, red circle indicates building flaw due to RBV, b) Horizontal view of machined sample, c) Vertical view of machined sampleCompression tests were performed according to the ASTM E9-19 Standard Test Methods of Compression Testing of Metallic Materials at Room Temperature. The eight samples were tested on an Instron 5985 universal testing machine with a 250 kN load cell. Samples were preloaded to 90 N and the strain rate used was 10(cid:237)3 min(cid:237)1.

The eight samples were tested on an Instron 5985 universal testing machine with a 250 kN load cell. Samples were preloaded to 90 N and the strain rate used was 10(cid:237)3 min(cid:237)1. The testing was completed at 25% strain since only the elastic behavior was of interest for this loading. Load and displacement were recorded, and calculations were completed after testing to develop stress-strain graphs and to record the 0.2% offset yield strength.2.3 Specifications and procedure for bend testingThe bending samples were fabricated in two build orientations (x and z) due to height limitations of the RBV chamber. After being removed from the build plate the parts were machined using a shaper to create equal dimensions and minimize surface roughness. Figure 5 shows the builds before (a, b) and after machining (c) and include the dimensions of the printed and tested specimen. There were four samples tested: three built in the x direction, one built in the z direction.

There were four samples tested: three built in the x direction, one built in the z direction. Similar to the compression sample build, the second z direction sample had clear defects, labeled in Figure 5 with the red circle, and was unable to be tested.21818.4 mm I60mm18.3 mmFigure 5: Bending samples before and after machining with labeled dimensions, a) build orientation x (horizontal), b) build orientation z (vertical), c) machined samples, red circle labels defected part of sampleThe width to thickness ratio of the samples are just over 2:1. The samples were tested using a 3-point bend fixture on an Instron 5985 universal testing machine with a 250 kN load cell. Figure 6 shows the sample mounted for testing as well as the orientation of the build direction in respect to the applied load.

Figure 6 shows the sample mounted for testing as well as the orientation of the build direction in respect to the applied load. The z build orientation refers to a build direction being parallel to the direction of the load, while the x build orientation refers to the direction of the build and the load being perpendicular to one another.Build directionXzFigure 6: Bend testing fixture with mounted sample, the build direction is labeled with axisThe samples were preloaded to 50 N. The deflection rate was 0.35 mm/min (strain rate of 1.417 x 10-4 min-1) and the samples were tested until failure. A span of 50 mm is used to allow for sufficient amount of the sample to extend past the end of the fixture.21822.4 Specifications and procedure for mini-tensile testingMini tensile samples were fabricated as rectangular blocks. After being removed from the build plate they were then EDM into mini-tensile dog bone samples.

After being removed from the build plate they were then EDM into mini-tensile dog bone samples. There were three blocks and five dog bone samples machined out of each block. Figure 7 shows the machined mini-tensile samples as well as a schematic of their dimensions. The samples were polished before testing by wet sanding using 600 grit. It should be noted that the top and bottom tension samples (labeled in Figure 7) were not used for testing due to their clear surface defects. The thickness of the samples is 1 mm and is not labeled in the dimension image. A more detailed description of the specifications of the mini tensile test can be found in the study done by Rios et al [10].BottomTopbc..,a.!. ::0 l 'Figure 7: a) machined mini tensile samples, b) dimensions of the samplesNine samples were tested in total for the mini-tensile testing. Three from each fabricated block. The samples were tested using a mini-tensile fixture on an Instron 5960 with a load cell of 10 kN.

Three from each fabricated block. The samples were tested using a mini-tensile fixture on an Instron 5960 with a load cell of 10 kN. The fixture set up can be seen in Figure 8. Samples were preloaded to 50 N and tested at a strain rate of 0.015 min-1. An extensometer can be seen on the fixture in Figure 8 and was removed during testing once the strain of the test reached 0.7%.Figure 8: Mini-tensile fixture setup, arrows indicating the use of the extensometer2183:-0 t'.l3. RESULTS AND DISCUSSION3.1 Compression testing resultsAll the compression tests were successfully performed for the eight samples consisting of the three different build directions. Table 2 shows the average 0.2% offset yield strength of each build direction. The z (vertical) build direction is shown to have the highest yield strength compared to the x and y (horizontal) directions. The x and y samples had very similar values.

The z (vertical) build direction is shown to have the highest yield strength compared to the x and y (horizontal) directions. The x and y samples had very similar values. The z direction was found to have an increased yield strength of over 10% when compared to the other two builds. The compressive yield strength from a previous study [3] was seen to be slightly higher, but the results are in good correlation with each other. The stress-strain curves for the samples are shown in Figure 9. The curves representing each build direction are the typical curves from the test. It is clear from curves on Figure 9 the samples built in the z direction have a higher yield strength and a slightly higher stress value throughout the entire test. Each orientation shows elastic behavior until about 2% strain before starting to plastically deform. This elastic area is to the left of the vertical green line, while the plastic region is to the right.

This elastic area is to the left of the vertical green line, while the plastic region is to the right. This vertical green line is also intersecting the curves right around the value of the samples’ yield strengths.Table 3: Compression testing 0.2% offset yield strength resultsBuild Direction Yield Strength (MPa)X Y Z252.2 ± 6.66 255.8 ± 2.91 299.2 ± 0.75Literature [3]Z317 ± 2Stress vs Strain600500..--..400 C'O a..~ 7:300(/) ~ en 200100o~~ -~ - - -~ - - -~ - - -~ - - -~ 0.25 00.050.15 0.1 Strain (mm/mm)0.2Figure 9: Stress-strain curves for solid sample compression testing, curves show typical results21843.2 Bend testing resultsThe four bending samples were successfully tested for their flexure strength using the 3-point bend fixture. The 0.2% offset flexure yield strength was calculated from the data curves made from the Instron universal testing machine.

The 0.2% offset flexure yield strength was calculated from the data curves made from the Instron universal testing machine. Table 3 shows the average yield strength to be 446.0 MPa for x build direction samples, and 448.4 MPa for the z direction samples. There were three samples for x direction which lead to a standard deviation of 8.09 MPa. The stress-strain curves are shown in Figure 10. The ‘x’ curve is a representative of the typical response of the samples while the ‘z’ curve is from the only tested sample. Similar to the compression graphs, Figure 10 shows that the samples show elastic behavior till about 1.2% strain. After 1.2% strain, the slope of the graphs starts to decrease to show plastic deformation. These two areas are separated by a green line. The elastic region is to the left while the plastic region is to the right.

These two areas are separated by a green line. The elastic region is to the left while the plastic region is to the right. Although only one z direction build was tested there was no clear effect of build orientation on the flexure yield strength.Table 4: Flexure yield strength due to 3-point bend testingBuild Direction X ZFlexure Yield Strength (MPa) 446.0 ± 8.09 448.4800Stress vs Strain700600ro a...5002 ';; 400V) ~ u5300200~ L--=:J100o .._- - -~ - - - - - - -L - -~ - - - -~ - - - -~ - - -~ 0.0500.010.02 0.03 Strain (mm/mm)0.04Figure 10: Stress-strain curve for solid sample bend testing, curves show typical results for the build orientation. The green line is indicating the flexure yield strength as well as the elastic and plastic behavior regions of the curves21853.3 Mini-tensile testing resultsFifteen mini-tensile samples were machined and nine were tested to complete fracture.

Table 4 shows the average yield strength of the nine tests to be 284.5 MPa with a standard deviation of just under 8 MPa. The stress-strain curves of two representative samples are shown in Figure 11. The vertical red line is the indicator for where the samples’ behavior went from elastic to plastic. In the plastic region of the graph there are dips in the stress values indicated by a red circle. This is due to the removal of the extensometer. The final dips in the graphs show the parts had sudden fractures. The average yield strength of the samples is slightly higher than the expected values from previous literature [3] but is in good agreement.Table 5: 0.2% offset yield strength due to mini-tensile testingBuild Direction Yield Strength (MPa)X&Y284.5 ± 7.56Literature [3]Z268 ± 2450Stress vs Strain400350ci,300 a..

~ ~250 (/J ~ in 200150100so ~-~--~-~--~--~-~--~-~--~ 0.180.020.040.060.080.10.120.140.160Strain (mm/mm)Figure 11: Stress-strain curve for solid sample mini-tensile testing, curves show typical results for the build orientation, the red line indicates the separation of elastic and plastic behavior regions, the red circle indicates the dip in stress caused by the removal of the extensometer4. CONCLUSIONSSelective laser melting was used in the RBV chamber of a Renishaw AM 250 to manufacture samples for compression, bending, and mini-tensile testing. The build orientation was varied for the compression and bending tests. The 0.2% offset yield strength and stress-strain curves have been calculated and reported for each type of loading. The compression test results showed that2186the samples built in the z direction had an increased yield strength of over 10% when compared to the sample built in the x and y directions.

The compression test results showed that2186the samples built in the z direction had an increased yield strength of over 10% when compared to the sample built in the x and y directions. The compressive yield strength from a previous study was found to be comparable. The bending test showed no clear difference in yield strength or behavior of the samples built in the x and z direction. The mini tensile samples were built in the x and y direction and were shown to have a yield strength that was comparable to previously reported tensile results. This showed our test samples built using the RBV of the Renishaw were on par with samples built using the entire build chamber.5. ACKNOWLEDGEMENTSThis research is sponsored by the Industrial Consortium of the Center for Aerospace Manufacturing Technologies (CAMT) at Missouri University of Science and Technology. The authors would also like to thank Honeywell Federal Manufacturing and Technology for the use of the Renishaw machine.6.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCompressive properties optimization of a bio-inspired lightweightstructure fabricated via selective laser meltingLiang Meng, Zemin Wang(cid:13), Xiaoyan Zeng Wuhan National Laboratory for Optoelectronics, Huazhong University of Science and Technology, Wuhan 430074, PR China *Corresponding author. Tel: +86-27-87544774; zmwang@hust.edu.cnAbstractCompressive properties optimization of a bio-inspired lightweight structure is developed by Response Surface Methodology (RSM) and Non-dominated Sorting Genetic Algorithm II (NSGA-II). Multi-layered bio-inspired structures of a Ti6Al4V alloy are designed and fabricated by Selective Laser Melting. The results show that the optimized structure parameters of bio-inspired structures can be obtained by RSM and NSGA-II.

The results show that the optimized structure parameters of bio-inspired structures can be obtained by RSM and NSGA-II. The relative error rate of experimental results and response values is less than 10%. Moreover, increasing the number of layers cannot effectively improve energy absorption (EA) and specific energy absorption (SEA) for multi-layered bio-inspired structures. The damage process of bio-inspired structures with different core-arranged configurations fails layer by layer. The load-displacement curves and damage process of FE simulations are consistent with the experimental results.Keywords: Bio-inspired structure; Selective laser melting; NSGA-II; Multi-layersIntroductionOwing to their high specific strength, anti-vibration, energy absorption, and excellent buffering capacity(cid:712)lightweight sandwich structures are widely used in aerospace, automobile and military applications [1-5].

However, higher requirements for excellent mechanical properties and lightweight structures are put forward with the rapid development of the aerospace and aviation fields [6]. However, nature has developed a mass of perfect performance structures and materials by evolving for thousands of years [7-9].Among them, the beetle elytra are considered as a better bio-inspired prototype owing to its particular microstructure and excellent mechanical properties [10]. Thus, inspired by the beetle elytra, a large number of lightweight structures have been designed, fabricated and analyzed [2, 11-12]. However, those investigations also indicated that bio-inspired structures were generally designed according to the shape or profile features of biological prototypes [13-15]. Excellent mechanical properties of bio-inspired structures are closely related to those structural parameters and their interactions though it is difficult to analyze [16].

Excellent mechanical properties of bio-inspired structures are closely related to those structural parameters and their interactions though it is difficult to analyze [16]. Thus, the structural optimization plays an essential role in understanding the relationship between structural parameters and mechanical properties. However, traditional optimization methods such as2280orthogonal design, homogeneous design, and factorial analysis methods not only take a number of experiments but also cannot explain the interactions between structural parameters well [17]. Comparatively, response surface methodology (RSM) [18] is considered as the most relevantly utilized method for optimization analysis, and it is also known for building empirical models and analyzing the influence of independent variables to several dependent variables [19].

Most importantly, RSM can obtain a specific mathematical equation which considers the influence of interactions between structural parameters on optimization variables. Moreover, the multi-objective optimization is very necessary in order to design a multi-functional bio-inspired lightweight structure. Genetic algorithm (GA) is a promising tool for solving multi-objective optimization problems [20], which has a good global search ability, fast process speed and high optimization precision when objective functions have specific equations. Thus, an important method is developed on the multi-objective optimization of bio-inspired structures by combining RSM with GA. The multi-objective optimization functions in GA are obtained by RSM, and there are few investigations on the multi-objective optimization of bio-inspired sandwich structures [21-24].In addition to the structural parameters, the layer numbers of cores may have a great influence on the mechanical properties.

A few references indicated that multilayer structures can improve the compressive performances of sandwich structures under out-of-plane loading [25-26]. However, most studies were mainly focused on single-layered sandwich structures [27-30]. The main reason is that the multilayer bio-inspired sandwich structures are difficult to fabricate by conventional processing methods due to their complex geometric features [21]. Selective Laser Melting (SLM) is a typical powder additive manufacturing process, which has higher dimensional precision and the capability to build any complex shapes of metal parts that would otherwise be difficult or impossible to be produced by using conventional manufacturing processes [31-32].

Thus, SLM is a practical method to manufacture the multilayered bio-inspired sandwich structures.In this paper, based on the bio-inspired sandwich structure designed by Meng [28], response surface methodology (RSM) and non-dominated sorting genetic algorithm II (NSGA-II) are used for multi-objective optimization in order to obtain lightweight structures with excellent comprehensive performances. Moreover, bio-inspired lightweight structures fabricated by SLM using optimal structural parameters are adapted to identify the accuracy of optimization models. Considering the influences of the layer number, multilayer bio-inspired sandwich structures are designed, fabricated by SLM and analyzed in detail under out-of-plane loading. Moreover, the damage process of bio-inspired structures is also investigated by experiments and FEM.Multi-objective optimization methodBased on the microstructure of Cybister elytra [28], the lightweight sandwich structure is designed, as shown in Fig. 1.

1. Response surface methodology (RSM) is2281processed in the Design Expert software which is used to optimize the structural parameters in this paper. RSM consists of a set of mathematical and statistical techniques to develop a functional relationship between a response of interest, y, and a number of associated control variables, (cid:1876)(cid:2869), (cid:1876)(cid:2870),…, (cid:1876)(cid:3038)[33]. Three structural variables consist of the core height h, the thickness of the core (cid:1872)(cid:2868) and thickness of panel (cid:1872) in this paper.Fig. 1. The geometric model of (a) bio-inspired structure (b) a unitThe response value y1 is Energy Absorption (EA), the dependency of response with process variables, h, t, t0, is illustrated in Eq.

1. The geometric model of (a) bio-inspired structure (b) a unitThe response value y1 is Energy Absorption (EA), the dependency of response with process variables, h, t, t0, is illustrated in Eq. (1).(cid:1877)(cid:2869) = (cid:1858)(cid:2869)((cid:1860), (cid:1872), (cid:1872)(cid:2868)) (1) In our work, the relationship between the values of the explanatory variables and the levels of parameterization considered is shown in Table 1. Moreover, the Box-Behnken Design method [34] is used in the software and a third order polynomial is adopted to obtain the appropriate model.

It can be seen that the relative errors of most tests are less than 10%, which also indicates that the predicted values are well consistent with the experimental values and the optimization model can explain the relationship between the variables and the response well.Table 3 Comparison of experimental and predicted values of energy absorption.TestExperimental value (J)Theoretical values (J)Relative error (%)14.343.929.67228326.299.0330.343 4 5 6 7 8 9 10 11 12 13 14 15 16 1742.00 41.00 30.00 49.12 36.16 33.48 53.36 16.28 26.49 13.85 10.32 44.12 51.90 111.20 5.0644.67 44.67 27.26 44.67 44.67 35.38 44.67 19.02 23.75 13.43 8.43 46.01 50.01 110.78 4.645.97 8.27 7.76 9.06 19.05 5.37 16.28 14.40 9.81 3.03 18.31 4.28 3.64 0.37 8.30Considering the lightweight and crashworthiness of bio-inspired structures, non-dominated sorting genetic algorithm II (NSGA-II) is used to achieve the multi-objective optimization [36].

3.Four layersThree layersTwo layersFig. 3. Multi-layered bio-inspired structures model2285Experimental proceduresGas atomized Ti-6Al-4V powders were used as the raw material in SLM experiments. Bio-inspired structures were fabricated in a self-developed SLM system (LSNF-1) [37]. In order to ensure the fabrication quality of bio-inspired structures, the optimal processing parameters were used to build the sandwich structures [28].Bio-inspired structures used for response surface analysis are shown in Fig. 4(a). Fig. 4(b) shows the multi-layered bio-inspired structures fabricated by SLM. It should be noted that the single layer sandwich structure was used to verify the multi-objective optimization model. The electronic universal testing machine (Shimadzu AG-100KN) was used for the flatwise compression tests.

The electronic universal testing machine (Shimadzu AG-100KN) was used for the flatwise compression tests. The indenter impacted the multi-layered bio-inspired structures at a velocity of v=1 mm/min during flatwise compressive tests and a camera recorded the entire process.EA and SEA [38, 39] are used to judge energy absorption capacity, and they arecalculated as follows.The total energy absorption (EA) is the area under the load-displacement curve:(cid:3031) (cid:1831)(cid:1827) = (cid:1516) (cid:1832)((cid:1876))(cid:1856)(cid:1876) (cid:2868)Where d is the displacement. It is generally believed that energy is absorbed before the structure is crushed, and F(x) is the magnitude of the load.The specific energy absorption is the absorbed energy per unit mass (SEA):(cid:1845)(cid:1831)(cid:1827) =(cid:1831)(cid:1827) (cid:1875)(11)Where w is the weight of whole bio-inspired structures.(a)(b).,,....,, ...# 17Fig.4.

The bio-inspired structure samples of multi-objective optimization (a), multilayer bio-inspired structures fabricated by SLM (b)Results and discussion4.1. Verification of the multi-objective optimization modelComparisons between experiment values of parameters and multi-objective optimization results are summarized in Table 4. Results reveal that the experiment value of EA (87.12 J) and density (1.52g/cm3) are slightly higher than the multi-objective optimization value of EA (81.56 J) and density (1.46g/cm3). The reason is that a certain manufacturing error exists in structural dimensions during SLM [40].

It shows the influences of the core height, thickness of panel and core on the energy absorption. It can be seen that the high thickness of core and panel will favor the high EA (Fig. 5(a) and (c)). Additionally, it seems that a proper core height, around 6 mm, can absorb more energy (Fig. 5(b)). Results also indicate that the interactions of structural parameters have a great impact on the compressive performances. Compared with core height, the interaction of the thickness of the panel and core has the greatest influence on the compressive properties.The th ickn ess of panel= 0.54 mmThe th ickness o f core = 0.80 mmFig. 5. Response surface with the core height and thickness of core (a), core height and the thickness of panel (b), the thickness of panel and core (c).4.3. Compressive properties of multi-layered bio-inspired sandwich structuresForce-displacement curves of multi-layered structures with parallel-arranged configurations are shown in Fig. 6.

Compressive properties of multi-layered bio-inspired sandwich structuresForce-displacement curves of multi-layered structures with parallel-arranged configurations are shown in Fig. 6. Generally, the compressive force-displacement curves for sandwich structures consist of three zones: elastic, buffer and crushing zone. Bio-inspired structures are not damaged in the elastic zone, and most energy is absorbed in the buffer zone. In addition, it is summarized that bio-inspired structures will fail when the force-displacement curves fall in the crushing zone [28]. From Fig.6, it should be noted that the sharp drop in force indicates that the multi-layered structures have been fractured. Many peaks appear in the force-displacement curves, which means that several areas of bio-inspired structures have been damaged under the crushing loading. It can be seen that core breaking is the main damage mode during the compression.

6. Force-displacement experimental curves of bio-inspired structure with parallel-arranged configurations: (a) single layer, (b) two layers, (c) three layers and four layers (d). curves almost had no peaks. The reason may be related to the intrinsic properties of materials. It should be noted that core breaking is the core material occurs to be yield and fractured when the material stress reaches to maximum stress, which belongs to material failure. Comparatively, the core buckling is the structures occur to be instability before the material stress reaches to maximum stress, which belongs to structural failure. Thus, the compressive force of core breaking is greater than that of core buckling. It also seems that the first peak forces are all about 20 kN with the increase of layer numbers, which indicates that increasing the layer number will have no influence on the first peak force. However, the maximum peak force firstly decreases to the first peak force and then keeps constant.

However, the maximum peak force firstly decreases to the first peak force and then keeps constant. It is because that the stability of whole bio-inspired structures decreases when the layer number increases. The maximum peak force is equal to the first peak force in the curves when the layer number is two. With the further increase of the layer number, the maximum peak force will keep constant and be equal to the first peak force.The damage photographic images of two-layered bio-inspired structures with parallel-arranged configurations are shown in Fig. 7. Under the compressive loading, the bottom layer of multilayer bio-inspired structures is firstly damaged and crushed. Then, the middle panel occurs to be fractured. Next, the upper layer of multilayer bio-inspired structures also occurs to be crushed owing to that the core has been fractured in several regions. Finally, the whole bio-inspired structure is compacted completely.

Finally, the whole bio-inspired structure is compacted completely. During compression, all cores and panels have been damaged. It is2288summarized that the core fracture is the main failure mode of bio-inspired structures under the compressive loading. The damage pictures also showed that several areas of core structures have been fractured, which also explains why the force-displacement curves have many peak forces.Energy absorption properties of multi-layered structures are shown in Fig. 8. It can be seen that with the increase of the layer number, the bio-inspired structures can absorb the same amount of energy (~ 80 J). The reason is that the force has dropped drastically even though the displacement increases. The SEA decreases gradually due to the weight of bio-inspired structures increases.

8. Energy absorption and specific energy absorption of multi-layered structures with parallel-arranged configurations4.4. The simulation results of multi-layered sandwich structuresThe ABAQUS software was used to simulate compression tests of bio-inspired sandwich structures, and the C3D8R element is selected. And the CAD model is the2289same as that of the bio-inspired structure. And the simulation model is shown in Fig. 9(a). The pressing plate is set as a rigid body. The contacts between the rigid body and structure are set as the surface to surface contact. Ti6Al4V alloy acts as the simulation material, whose physical parameters are given in Table 4 and the stress-strain curve is shown in Fig. 9(b).12001000- SLMed Ti6Al4V2000.010.020.030.04 Striai10.050.060.07Fig. 9.

9(b).12001000- SLMed Ti6Al4V2000.010.020.030.04 Striai10.050.060.07Fig. 9. The simulation model of compressive tests (a) and the stress-strain curve of Ti6Al4V (b)Table 4 The physical parameters of Ti6Al4V alloyParameters Density (g/cm3)Elastic modulus (MPa)Poisson’s ratioValue1.521100000.34The FEM and experimental results of multi-layered bio-inspired sandwich structures are shown in Fig. 10. It can be seen that the load-displacement curves of FEM result are consistent with that of experimental results basically. The difference is that the load-displacement curves of compression experiments have many zones which their load is close to zero. The reason is that the Ti6Al4V alloy was brittle fractured on the compressive tests and the velocity of the compressive head is 1mm/min. Therefore, the load will decrease rapidly to zero when the sandwich structures break.

Therefore, the load will decrease rapidly to zero when the sandwich structures break. However, in the FE simulation, the whole compressive time is set as 0.1s, and the time is so short that the load cannot decrease to zero. The damage process of two-layered sandwich structures is shown in Fig. 11, and the results indicated that the FEM is consistent with the experiments. Moreover, the fracture zones and stress concentration locations in FEM also are consistent with the experimental results, which declare that the finite element model is accurate.22900.0810030(a)80- Experimcantal -- FEM(b)25- Experimental - - FEMz60 .,, C, "' 0 ...l 4020zC, 81 5 :s"--1020054 Displacement (mm)310 8 Displacement (mm)61214(c) 30- Experimental - - FEM(d) 25- Exper imen tal - - FEM2520203~15 ~ 0 "--~ISu e0 "-101046101214Displccmcnt (mm)16182022101214Displacement (mm)1618202224Fig. 10.

10. The FEM and experimental results of sandwich structures with a single layer (a), two-layered (b), three-layered (c) and four-layered (d) core.(a)(b)I] ' EBt I(d)I] ''"Fig. 11. The damage process of two-layered sandwich structures: D=0mm (a), D=1mm (b), D=1.50mm (c) and D=8.50mm (d).ConclusionsThe compressive properties optimization of bio-inspired structures via RSM and NSGA-II has been carried out in this work. Moreover, multi-layered bio-inspired structures of Ti–6Al–4V alloy are designed, fabricated by SLM and analyzed in detail. The conclusions are as follows:(1) Response surface methodology (RSM) and non-dominated sorting genetic algorithm II (NSGA-II) can be applied to better optimize the compressive properties of bio-inspired structures fabricated by SLM.2291the EA and SEA of multi-layered bio-inspired structures with parallel-arranged configurations.

The damage process of bio-inspired structures with different core-arranged configurations fails layer by layer.(2) Increasingthelayer number cannotimprove(3) The load-displacement curves and damage process of FE simulations areconsistent with the experimental results.AcknowledgmentsThis work is supported by the Fundamental Research Funds for the Central Universities (HUST: 2016YXZD005). Special thanks for the Analytical and Testing Center and the State Key Laboratory of Materials Processing and Die & Mould Technology of HUST.References[1] L. Murr, S. Quinones, S. Gaytan, M. Lopez, A. Rodela, E. Martinez, D. Hernandez, E. Martinez, F. Medina, R. Wicker, Microstructure and mechanical behavior of Ti–6Al–4V produced by rapid-layer manufacturing, for biomedical applications, J. Mech. Behav. Biomed. mater. 2 (2009) 20-32.[2] C. Guo, D. Li, Z. Lu, Mechanical properties of a novel, lightweight structureinspired by beetle elytra, Chin. Sci. Bull.

Phone: 8067423543. Fax: 8067423411 Email address: weilong.cong@ttu.eduAbstractTitanium (Ti) and its alloys are widely used in chemical, astronautical, and biomedical industries. However, their poor load endurance properties affect their fields of applications especially under severe loading conditions. To enhance these properties, TiBw reinforcement was synthesized by in situ chemical reaction between elemental Ti and boron. Strong interfacial bonding between TiBw reinforcement and Ti matrix was obtained due to the in situ chemical reaction. Owing to its capability of producing difficult-to-machine bulk composites with uniform properties, laser engineered net shaping (LENS) technique was utilized to fabricate TiBw reinforced Ti matrix bulk composites. Few researches have been reported on these three- dimensional metal based bulk composites by using LENS. In this work, effects of TiBw reinforcement and laser power on compressive properties were investigated.

In this work, effects of TiBw reinforcement and laser power on compressive properties were investigated. The microstructures of the fabricated parts were observed and analyzed by using scanning electron microscopy.KeywordsLaser engineered net shaping (LENS); TiBw reinforcement; Titanium matrix composites (TMCs); Compressive properties1 IntroductionAs excellent corrosion resistance metallic materials, titanium (Ti) and its alloys are recognized for their high strength-to-density ratio and outstanding biocompatibility, thus rendering wide applications in chemical, astronautical, and biomedical industries. [1,2]. However, for severe working conditions where high bearing capacity property is necessary (e.g. tooth’s root implants [3], actuator strut in the engine thrust direction control system [4], etc.), ceramic reinforcements are required to improve compressive properties of Ti and its alloys.

tooth’s root implants [3], actuator strut in the engine thrust direction control system [4], etc.), ceramic reinforcements are required to improve compressive properties of Ti and its alloys. Compared with other ceramic reinforced Ti matrix composites (TMCs), titanium boride whisker (TiBw) reinforced TMCs (TiBw-TMCs), exhibiting more benefits, are of particular interest. (1) TiBw reinforcement is chemically compatible with TMCs since there are no intermediate phases between TiB and Ti phases [5]. (2) The residual stresses and stress concentration effects induced by TiBw reinforcement can be weakened or eliminated as a result of comparable thermal expansion coefficients and densities between TiB and Ti [6].

(3) Strong interfacial bonding1 512between TiB and Ti can be formed as a result of excellent thermodynamic and chemical stability between these two phases [7,8].In recent years, in situ TMCs have gained their popularity due to the benefits exhibited by in situ TMCs over ex situ TMCs [9]. Strong interfacial bonding between TiB and Ti as well as better mechanical properties can be obtained with the formation of finer TiBw and the uniform distribution of rigid ceramic TiBw reinforcement in the Ti matrix [9-11]. In addition, the energy generated during chemical reaction between elemental Ti and boron (B) can be utilized to facilitate fabrication process.Conventional fabrication processes (e.g. casting, powder metallurgy, etc.) of ceramic reinforced TMCs are usually time consuming and labor intensive [8,12,13]. In addition, post- processing, which requires new tooling and production line, is needed to assist fabricating complex-shape parts.

In addition, post- processing, which requires new tooling and production line, is needed to assist fabricating complex-shape parts. To overcome these problems, laser additive manufacturing (LAM), the concept of which is free-to-design, is proposed to fabricate ceramic reinforced TMCs [8,10]. Among all the LAM techniques, laser engineered net shaping (LENS) has specific advantages of small heat affected zone, easy-to-control cooling, small substrate deformation, and changeable dilution [14]. In addition, the unique features of LENS make it an irreplaceable technique in part repair of worn components [15]. Therefore, LENS has been used to fabricate bulk parts with mixed Ti and B powders in this paper.Figure 1 shows the illustration of LENS process. The powder stream generated from the powder feeder was melted by the laser beam ejected from the laser source, forming a high- temperature molten pool. The molten pool would catch more powders to increase its volume.

The molten pool would catch more powders to increase its volume. When the laser beam moved away, the molten pool solidified due to heat dissipation. Following the trajectory designed by the CAD file, one layer was built with the moving of deposit head. Afterwards, the deposit head ascended, leaving space for the next layer fabrication. In the end, the final bulk part was built by repeating these procedures layer by layer.Scanning directionLaser sourceLaser mirrorLaser beamFocus lensShield gasDeposit headZPowder streamMolten poolDeposited layerSubstrateFigure 1. Illustration of LENS process2 513The remainder of this paper is organized as follows: Section 2 presents descriptions of experiments and measurements procedures. Effects of TiBw reinforcement and laser power on the compressive properties of the fabricated bulk parts will be analyzed and discussed in Section 3.

Effects of TiBw reinforcement and laser power on the compressive properties of the fabricated bulk parts will be analyzed and discussed in Section 3. Section 4 will describe the conclusions drawn from this paper.2 Experiments and measurements procedures2.1. Powder treatmentCommercially pure Ti (CP-Ti) powders (Atlantic Equipment Engineers Inc., Upper Saddle River, NJ, USA) (average size of 150 μm, purity of 99.7%) and B powders (Chemsavers, Inc., Bluefield, WV, USA) (average size of less than 5 μm, purity of more than 96%) were used in this paper. According to the Ti-B phase diagram, eutectic solidification occurred, when the weight percent of B was around 1.6%, with the formation of nano-size TiB whiskers, thus rendering refinement of Ti grains and better mechanical properties [10,16]. It was reported by Wei et al. that high fracture toughness and bending strength were achieved at this weight ratio for TiBw-TMCs [17].

It was reported by Wei et al. that high fracture toughness and bending strength were achieved at this weight ratio for TiBw-TMCs [17]. In this paper, the CP-Ti powders were premixed with 1.6 wt.% B powders by planetary ball mill machine (ND2L, Torrey Hills Technologies LLC., San Diego, CA, USA). Based on preliminary results, the ball mill parameters were set at optimized level: including rotation speed of 200 rpm, ball-to-powder weight ratio of 5:1, and milling time of four hours. After ball mill treatment, flowability and compactability of the powders were improved. In addition, the B powders were well distributed in the Ti powders.2.2. Experimental conditionsTable 1 shows process parameters adopted in this paper. The input fabrication variables, such as scanning speed (11 mm/s), powder feed rate (1.65 g/min), hatch distance (381 μm), etc., were optimized based on preliminary results.

The input fabrication variables, such as scanning speed (11 mm/s), powder feed rate (1.65 g/min), hatch distance (381 μm), etc., were optimized based on preliminary results. The scanning direction, initiating from 45o, rotated 90o for each layer to weaken the effects of scanning orientation on the fabricated parts. In order to prevent oxygen from contaminating fabricated parts, the oxygen level was kept below 500 ppm. The dimensions of the fabricated cylindrical parts were ϕ6 mm15 layers.ParametersTable 1. Process parametersValues or range× UnitsScanning speed11mm/sPowder feed rate1.65g/minHatch distance381μmNumber of layersScanning orientation1545, 135, 225, and 315/ oOxygen level<500ppm2.3. Measurement proceduresThe fabricated parts were cut through cross-section, then ground and polished by a grinder-polisher machine (MetaServ 250, 49-10055, Buehler, Lake Bluff, IL, USA).

Measurement proceduresThe fabricated parts were cut through cross-section, then ground and polished by a grinder-polisher machine (MetaServ 250, 49-10055, Buehler, Lake Bluff, IL, USA). Afterwards, these parts were etched with Kroll’s reagent (HF: 3%; HNO3: 6%; water: balance) for ten3 514seconds. To observe and analyze the microstructural characterizations of the parts, the scanning electron microscope (SEM) (S-4300, Hitachi, Tokyo, Japan) was used.Before compressive test, both planar sides of each part were ground and polished by the grinder-polisher machine. Along the deposit head moving direction (Z), as is shown in Figure 1, compressive test was performed on the fabricated cylindrical parts by using a tensile tester (AGS-50kNXD, Shimadzu, Kyoto, Kyoto Prefecture, Japan), at a constant cross-head speed of 0.005 mm/s.

For each condition, the tests were conducted at least three times at room temperature to have better statistics.3 Experiment result and discussionFigures 2(a) and 2(b) show effects of laser power on the defects (e.g. micro-cracks and micro-pores) formation of fabricated CP-Ti parts. It can be seen in Figure 2(a) that not only micro-pores but also micro-cracks presented. In Figure 2(b), micro-pores were dominant, whereas micro-cracks could hardly be seen. The reason lies in the fact that micro-cracks were prone to form at the interfaces of neighboring fabricated layers as a result of insufficient laser power (125 W) [14]. By increasing the laser power to 200 W, the top part of the former solidified layer could be remelted, resulting in a better interfacial bonding between neighboring layers with less cracks or pores. As is reported, the defects, from where dislocations easily initiated with applied force, were stress concentrators, resulting in weakened mechanical properties [18].

As is reported, the defects, from where dislocations easily initiated with applied force, were stress concentrators, resulting in weakened mechanical properties [18]. Figure 2(c) shows the cross-section characterizations of TiBw-TMCs processed with 200 W laser power. In LENS process, in situ exothermic chemical reaction between Ti and B occurred, forming prismatic TiBw reinforcement. As a kind of rigid ceramic reinforcement, TiBw contributed to the strengthening effect on the TiBw-TMCs [19]. In addition, the exhibition of TiBw reinforcement could lead to Ti grain refinement, which was also beneficial to the mechanical properties [6,10].(a)(b)(c)Micro-cracksMicro-poresMicro-poresTiB reinforcement300 μm300 μmFigure 2. SEM observations of cross-section characterizations of (a) CP-Ti at 125 W; (b) CP-Ti at 200 W; (c) TiBw-TMCs at 200 W.Compressive tests were performed to investigate effects of TiBw reinforcement and laser power on compressive properties.

As was reported, a compressive test was not as susceptible as a tensile test to minor defects. The minor defects inside the solid parts could cause premature failure and then affect property analysis [10]. Figure 3 shows effects of TiBw reinforcement and4 5156 μmlaser power on ultimate compressive strength (UCS) of the fabricated parts. The central tendency was represented by the average value of the measured data under each condition. The standard errors, which were calculated by dividing the standard deviation by the square root of the number of measurements, were used to represent error bars’ positive and negative values. It can be seen that the UCS was significantly increased with the increase of laser power from 125 W to 200 W for both CP-Ti parts and TiBw-TMCs. Higher densification degree with no obvious micro-cracks of the parts fabricated with 200 W laser power, as aforementioned, could be one possible reason.

Higher densification degree with no obvious micro-cracks of the parts fabricated with 200 W laser power, as aforementioned, could be one possible reason. Another possible reason could be larger residual stress induced by 200 W laser power, leading to the enhancement of UCS. Owing to the present of TiBw reinforcement, TiBw-TMCs exhibited superior UCS, as compared with that of CP-Ti parts at corresponding laser power level. It can be concluded that a dramatic increment of 38.2% (from 760.9 MPa to 1051.4 MPa) could be achieved by importing TiBw reinforcement into the Ti matrix as well as choosing high laser power (200 W). Similarly, yield strength increased with the increase of laser power for both CP- Ti parts and TiBw-TMCs, as is shown in Figure 4. At all levels of laser power, yield strengths of TiBw-TMCs were higher than those of CP-Ti parts.1100CP-Ti1000TiB-Ti) a P M900(S C U8005μ700125200Laser power (W)Figure 3.

Effects of TiBw reinforcement and laser power on ultimate compressive strength (UCS).1000CP-Ti) a P M850TiB-Ti(h g n e r t s d e Yt700l550i5μ400125200Laser power (W)Figure 4. Effects of TiBw reinforcement and laser power on yield strength.5 516Effects of TiBw reinforcement and laser power on maximum strain, which could be utilized to represent ductility, are shown in Figure 5. The present of TiBw reinforcement led to smaller average maximum strain than that of CP-Ti parts. In another word, the TiBw reinforcement was detrimental to the ductile properties of TiBw-TMCs. Huang et al. reported that ceramic reinforced TMCs exhibited low room-temperature damage tolerance (e.g. ductility and fracture toughness) due to the lack of ductile phase which could inhibit crack propagation [20]. For CP-Ti parts, the increase of laser power led to the decrease of the average maximum strain from 64.9% to 54.3%.

For CP-Ti parts, the increase of laser power led to the decrease of the average maximum strain from 64.9% to 54.3%. The high maximum strain value could possibly be ascribed to the high compatibility induced by the low densities of the parts fabricated with 125 W laser power. For TiBw-TMCs, a slight increment of maximum strain from 24.8% to 26.2% was achieved by increasing laser power from 125 W to 200 W. Figure 6 shows effects of TiBw reinforcement and laser power on toughness. It can be seen that the present of TiBw reinforcement had negative impacts on toughness. For both CP-Ti parts and TiBw-TMCs, toughness increased with the increase of laser power.70)%60( n a r t s m u m x a Mii5040CP-TiTiB-Ti305μ20125200Laser power (W)Figure 5. Effects of TiBw reinforcement and laser power on maximum strain.500) 3m / J 6 0 1 X400CP-Ti(300s s e n h g u o T200TiB-Ti5μ100125200Laser power (W)Figure 6.

Effects of TiBw reinforcement and laser power on maximum strain.500) 3m / J 6 0 1 X400CP-Ti(300s s e n h g u o T200TiB-Ti5μ100125200Laser power (W)Figure 6. Effects of TiBw reinforcement and laser power on toughness.6 5174 ConclusionsIn this investigation, bulk CP-Ti parts and TiBw-TMCs parts were fabricated and effects of TiBw reinforcement and laser power on compressive properties were investigated. The main conclusions are drawn as follows:(1) By increasing laser power from 125 W to 200 W, the densification degree of the parts fabricated could be improved with fewer defects inside. Due to this reason, the UCS and the yield strength increased with the increase of laser power. The increased laser power adversely affected maximum strain, whereas it positively affected toughness.(2) Prismatic TiBw was formed as a result of the chemical reaction between elemental Ti and B in LENS process.

Due to the grain refinement and strengthening effects induced by TiBw reinforcement, both the UCS and the yield strength were enhanced. The lacking in ductile phase of TiBw-TMCs led to weakened maximum strain, thus rendering worse toughness.References[1] Donachie, M. J. (2000). Titanium: a technical guide. ASM international. [2] Liu, X., Chu, P. K., & Ding, C. (2004). Surface modification of titanium, titanium alloys, and related materials for biomedical applications. Materials Science and Engineering: R: Reports, 47(3), 49-121. [3] Laoui, T., Santos, E., Osakada, K., Shiomi, M., Morita, M., Shaik, S. K., ... & Takahashi, M. (2006). Properties of titanium dental implant models made by laser processing. Proceedings of the Institution of Mechanical Engineers, Part C: Journal of Mechanical Engineering Science, 220(6), 857-863. [4] Spowart, J. E., & Clyne, T. W. (1999). The axial compressive failure of titanium reinforced with silicon carbide monofilaments.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCOMPRESSIVE RESPONSE OF STRUT-REINFORCED KAGOME WITH POLYURETHANE REINFORCEMENTRinoj Gautam1, Vijay Shankar Sridharan2, Wee Lee Teh3, and Sridhar Idapalapati1,2,31Singapore Centre for 3D Printing, School of Mechanical and Aerospace Engineering, Nanyang Technological University, 50 Nanyang Avenue, Singapore 6397982Rolls-Royce@NTU Corporate Lab, Nanyang Technological University, 50 Nanyang Avenue, Singapore 6397983School of Mechanical and Aerospace Engineering, Nanyang Technological University, 50 Nanyang Avenue, Singapore 639798AbstractLattice structures find immense application in lightweight structures for their high specific strength, modulus, and energy absorption. Strut-reinforced Kagome (SRK) structures provide better compressive performance compared to many existing lattice structures.

Strut-reinforced Kagome (SRK) structures provide better compressive performance compared to many existing lattice structures. In this study, the performance of acrylonitrile butadiene styrene (ABS) SRK lattice structures, fabricated by fused deposition modeling, under compression loading is investigated. Further, SRK structures were filled with different polyurethane in the empty space and their effect on the compressive performance was examined. The SRK structure demonstrated abrupt failure at the joints in the vicinity of face sheet, thereby reducing the energy absorption of the structure. The SRK with flexible foam (low-density polyurethane foam) had no significant effect on peak failure load and moduli, whereas energy absorption per unit mass was higher by 16.5%.

The SRK with flexible foam (low-density polyurethane foam) had no significant effect on peak failure load and moduli, whereas energy absorption per unit mass was higher by 16.5%. The SRK with the rigid foam (high-density foam) displayed not only the better energy absorption per unit mass (116%) but also different failure behavior than SRK only.IntroductionSandwich panels consist of lightweight core material in between two stiff and strong face sheets. The superior bending stiffness, high specific strength, and energy absorption of the sandwich structures give them a good potential in aerospace, marine and automobile industries [1]. The cellular core structures are either bending dominated or stretching dominated depending on their architecture [2]. Stochastic foam structures deform by the bending of the cell walls whereas the lattice structures deform either by the stretching or bending.

Stochastic foam structures deform by the bending of the cell walls whereas the lattice structures deform either by the stretching or bending. The stretch dominated structures have higher stiffness and strength than the bending dominated structures [3]. The low relative density lattice structure collapses either by elastic or plastic buckling and the structure with high relative density fail by yielding. The foam structures have superior energy absorption ability and commonly used in energy absorption applications [4, 5].Additive manufacturing is an innovative technology which made possible to fabricate different lattice structures with complex designs. The performance of Kagome structure fabricated2188through additive manufacturing under compression loadings has been investigated. Ullah et al.

The performance of Kagome structure fabricated2188through additive manufacturing under compression loadings has been investigated. Ullah et al. studied the performance of Kagome structure fabricated by selective laser melting and found that Kagome structures exhibited higher compressive strength than the conventional honeycomb structure [6]. The structures also showed similar effective modulus and energy absorption of conventional honeycomb structures [7]. Gautam et al. investigated the effect of part-built orientation and surface roughness on the compressive performance of Kagome structure fabricated by fused deposition modeling [8]. The graded relative density design of multi-layer Kagome structures exhibited higher energy absorption than the uniform density [9]. On their further studies, they studied the compressive performance of the modified Kagome structures.

On their further studies, they studied the compressive performance of the modified Kagome structures. For the same relative density, modified Kagome exhibited higher strength and stiffness than Kagome structures [10, 11].Different foams have been utilized as fillers in hollow cores to have better energy absorption of the sandwich panels. A study on the use of polyurethane foam on the circular cell honeycomb exhibited the stable collapse of the structure leading to the amplification of energy absorption [12]. Similarly, the composite egg box panels filled with the foam improved the energy absorption capacity [13]. Zhang et al. investigated the response of the pyramidal core structure filled with polyurethane foam under compression [14]. The foam has a synergetic effect on load carrying capacity and the structure filled with higher density foam exhibited superior energy absorption than unfilled structures. Kao et al.

The foam has a synergetic effect on load carrying capacity and the structure filled with higher density foam exhibited superior energy absorption than unfilled structures. Kao et al. studied the effect of foam on the impact performance of bi-material structures [15]. The inclusion of the foam structure with high ductility improved energy absorption due to excessive local and global deformation.In literature, we found that Kagome and SRK lattice structures demonstrated excellent strength and stiffness and foam structures exhibited tremendous energy absorption capacity. Hence, the study on the use of foam in the lattice structures under compression loading is needed. This study focuses on the use of polyurethane foam as the filler material to enhance the energy absorption capacity of the sandwich structure with strut-reinforced Kagome core. Strut reinforced structures were fabricated through fused deposition modeling with ABS material.

Strut reinforced structures were fabricated through fused deposition modeling with ABS material. Two different polyurethane foam with different densities were used as the fillers in the SRK core and their effect on the compressive performance of the structure was investigated.Experimental MethodsStrut-reinforced Kagome is an improvised version of Kagome with a vertical strut in the center as shown in Figure 1. The performance of the lattice structures depends on the relative density of the structure. Thus, unit SRK structure is fully defined by the geometric parameters: strut diameter (d), the inclination of the slant strut (θ) and the height of the core (h). In this study, the strut diameter was designed as 2.3 mm, the core height of 13 mm, the inclination angle of 54.74° and face sheet thickness was 2 mm. SRK core structures, with 5 x 5 units, were designed for the compression test of sandwich structure.

SRK core structures, with 5 x 5 units, were designed for the compression test of sandwich structure. The arrangement of the unit structures is illustrated in Figure 2.2189Figure 1: Unit SRK structure with its geometric parametersSide viewFront viewIsometric view Figure 2: Side, front and isometric views of the sandwich structure with SRK coreThe SRK structure samples were fabricated using ABS material through fused deposition modeling (FDM) on Stratasys® F370. The support material was removed with the help of detergent (Ecoworks™ cleaning agent) water. Two different (flexible and rigid) types of polyurethane foams were selected to fill the open space of the core. Flexible low-density polyurethane foam (FlexFoam-iT!™ III) and rigid high-density polyurethane foam (Foam-iT!™ 10) supplied by Smooth-On Inc. were used to fill the empty space in the core. The mold was designed to snugly fit SRK core sandwich and fabricated using the FDM.

were used to fill the empty space in the core. The mold was designed to snugly fit SRK core sandwich and fabricated using the FDM. The foam was filled in the mold-sandwich assembly such that the filling was homogenous and contained within the assembly.2190The compression test on all unfilled and foam filled SRK samples were carried out on Instron 5569 universal testing machine (UTM) with 50 kN load cells with the crosshead speed of 0.2 mm/min. An external linear variable differential transformer (LVDT) was used to record the displacement of cross head of the machine. The displacement recorded by LVDT and the force recorded by data acquisition system of UTM was used to calculate the compressive stress and strain of the sandwich structures, respectively.The properties of the foam materials were obtained by the compression test on the cube samples of 30 x 30 x 30 mm.

The compression test on low-density polyurethane foam samples was carried out on 500 N load cell whereas the test on high-density polyurethane foam was carried out on 50 kN load cell.Results and DiscussionThe stress-strain curves of both flexible foam and rigid foam polyurethane are shown in Figure 3. Rigid polyurethane foam demonstrates distinct linear region whereas it is difficult to notice the linear region in the flexible foam due to its inferior load carrying capacity in the low strain. The linear region is due to the elastic deformation of the foam which occur by the cell wall bending of the foam. The near-plateau region formed due to the loss of stability of the foam and compacting drastically with almost no response to the applied load. The plateau region is followed by the densification region where the cells of the foam completely flatten, and the load increases considerably with little deformation.

The plateau region is followed by the densification region where the cells of the foam completely flatten, and the load increases considerably with little deformation. The rapid increase in the stress with a small increase in the strain is due to the densification (contact of cellular surfaces with each other). The onset of the densification in stress-strain curve is the strain to densification. The rigid foam has high plateau stress and the strain to densification is less than that of flexible foam as shown in Figure 3.0.064cu -c. 0.04 ~ u, u, Q) Jj0.02 enregionci 3 a.~ - I.. ... en 1u, u, Q)2Densification regionPlateau reg ion0.000.00.20.40.6Strain(a)0.81.0i.o --Linear region0.20.4 Strain (b)0.6Figure 3: The compressive strain-strain curve for (a) flexible foam and (b) rigid foamThe average weight of the unfilled SRK samples, flexible foam (FF) filled, and rigid foam (RF) filled are listed in Table 1.

The increase in the weight of the SRK samples with the filling of flexible foam and rigid foam is 28.12% and 50.96%, respectively.21910.8Table 1: Average weight of the unfilled SRK and foam filled SRK samples.Average weight (g)Unfilled SRK 33.14 ± 0.07FF SRK 42.46 ± 0.11RF SRK 50.03 ± 0.21The representative stress-strain curve of SRK structure under compression is depicted in Figure 4. The plot of unfilled SRK and flex foam filled SRK are presented in Figure 4 (a) to see the contribution of the flexible foam on the performance. FF SRK structure exhibited similar compressive behavior as the unfilled structure. The stress-strain curve of both unfilled SRK and FF SRK show the linear region to the initial peak stress followed by the rapid decrease in the stress. The presence of FF did not have any influence in the initial peak strength. The load supported by FF is negligible and the truss core supported the whole load.

The presence of FF did not have any influence in the initial peak strength. The load supported by FF is negligible and the truss core supported the whole load. Once the truss started to fail, the structure did not support any load. The use of FF provided no support to the truss hence have the similar stress-strain curve as unfilled SRK.54 ns a..~ 3 -(/) (/) e 2 ..., en 1Unfilled SRK - - - - - Flex foam filled SRKI I I65 - - - - Rigid foam filled specimen ~ -(/)3 ' '5 - - - - Rigid foam filled specimen ~ -(/)3 , .. 'Unfilled specimenns a..4i.o0.2,-, , ,,' , 0.4 Strain (a)Q.6f.densi 0.81i.o0.20.4Strain (b)Figure 4: The compressive response of unfilled SRK and (a) flex foam filled SRK and (b) rigid foam filled SRKThe representative stress-strain curve of SRK and RF SRK under compression is shown in Figure 4(b). The use of rigid foam increased the initial peak strength of the structures by about 60%.

The use of rigid foam increased the initial peak strength of the structures by about 60%. This is because the RF has high density and they provided lateral support to the truss structure. After reaching the initial peak strength, the rigid foam filled displayed the plateau region. The stress continues to remain the constant, unlike the unfilled SRK structure. The cell walls of the foam began to collapse which progressed at the same constant load. The densification of the collapsed cell walls led to the densification of the structure. The densification of rigid foam filled structure occurred at lower strain than that of SRK structure.The energy absorption per unit volume of the structure is obtained by the area under the stress-strain curve. It can be observed that the area under the stress-strain curve of RF SRK is bigger than that of unfilled SRK.

It can be observed that the area under the stress-strain curve of RF SRK is bigger than that of unfilled SRK. The effect of the foam on the energy absorption of the structure can be compared with the energy absorption per unit weight of the structure, which considers the influence of added weight. The comparative performance of unfilled SRK and foam filled structure in terms of volumetric energy per unit mass is illustrated in Figure 5. There is a slight increase in2192' I0.6the energy absorption to densification with the flexible foam filling in SRK core. The SRK with rigid foam filling exhibited higher (116%) energy absorption than unfilled SRK structure.40 . - - - - - r - - -~ - - - - - - r - -~ - - , - - - - - - , 1.0ci ,,,~ 35 - •- Densification Strain .§ 30Energy absorbed__________ ,"")~ 25 'C 1l 20 ... 0 ~ 15 t'll ~ 10 ...

- - - - - r - - -~ - - - - - - r - -~ - - , - - - - - - , 1.0ci ,,,~ 35 - •- Densification Strain .§ 30Energy absorbed__________ ,"")~ 25 'C 1l 20 ... 0 ~ 15 t'll ~ 10 ... ! 5 w1-----0.8 C -~ -en0.6 C 0 ~(.)0.4 !Efl) C Cl) 0.2 CFigure 5: Total energy absorption to densification and densification strain of unfilled SRK, flex foam filled SRK and rigid foam filled SRK.The cumulative energy of absorption per unit mass of unfilled and foam filled SRK is plotted against the strain in Figure 6. The curve for the rigid foam filled SRK is different from that of foam filled SRK and unfilled SRK. The area for the initial region of foam filled SRK is less than the unfilled SRK. In initial loading, the truss supports the applied load in the foam filled structures. Thus, the foam filled SRK have less energy absorption in the initial linear region because of the increased weight of the structures.

Thus, the foam filled SRK have less energy absorption in the initial linear region because of the increased weight of the structures. The FF is very flexible and contributes negligibly in supporting the load, thus have similar energy absorption curve of unfilled SRK. The plateau region in the curve for FF SRK and unfilled SRK is due to the abrupt decrease in the load carrying capacity after the initial failure as shown in Figure 4(a). But the energy absorption curve of RF SRK increases almost linearly with strain. The plateau region of the stress-strain curve of RF SRK contributed to increasing the area under the curve, hence, increased the energy absorption. The rigid foam has provided the synergetic effect on the energy absorption capacity of the SRK structure.- Unfilled SRK · · · · · · Flex form filled SRK - - - - Rigid form filled SRK /,,'i.

The rigid foam has provided the synergetic effect on the energy absorption capacity of the SRK structure.- Unfilled SRK · · · · · · Flex form filled SRK - - - - Rigid form filled SRK /,,'i. "!-.,_~--..i.o0.20.4 Strain0.60.8Figure 6: The energy absorption per unit weight of unfilled SRK and foam filled SRK against the strain.2193ConclusionIn this study, the response of strut-reinforced Kagome core sandwich structure filled with different densities polyurethane foam under quasi-static compression loading was investigated. The compression test on unfilled SRK structure exhibited a rapid decrease in the load carrying capacity after reaching the initial peak strength. FF with SRK core could not resist the abrupt failure of the structure hence barely influenced the performance of the structure. The use of RF in the SRK core not only increased the peak load of the structure but constrained the rapid failure of the structure.

The use of RF in the SRK core not only increased the peak load of the structure but constrained the rapid failure of the structure. The RF provided the lateral support to the SRK truss increasing the peak load. The RF with SRK truss demonstrated the synergetic effect on the energy absorption capacity of the structure. The volumetric energy absorption per unit weight (considering the increase in the weight) increased by 116% with the use of RF in SRK core structure. The selection of appropriate density foam and its use on the lattice core sandwich structure can be utilized to increase the performance of the sandwich structure.AcknowledgmentsThis work is funded by Ministry of Education - Singapore (MoE) Academic Research Fund Tier 1 grant # MOE RG173/15 administered through Nanyang Technological University, Singapore.References[1] press1999. [2] N. Fleck, V.S. Deshpande, M.F.

Additive manufacturing has provided a unique opportunity to fabricate such complex geometries. In addition to the unit cell type, the strut size and shape can significantly affect the mechanical properties achieved. Therefore, furnishing a lattice structure library may help in selecting the appropriate combination of lattice types and dimensions for targeted mechanical performance for a specific application. This study presents a method for determination of mechanical properties, including strength and stiffness, for lattice structures. Finite element (FE) simulations are used as the main tool and the results of which are to be verified by mechanical testing of samples fabricated using the laser beam powder bed fusion (LB-PBF) process. Proper lattices with the stiffness matched with associated bone were determined. However, the result indicated that lattices made from 316L SS are not strong enough for bone implants.

Proper lattices with the stiffness matched with associated bone were determined. However, the result indicated that lattices made from 316L SS are not strong enough for bone implants. The proposed procedure can be used for other unit cells of interest due to its generality.IntroductionIn order to create biomedical implants that do not have negative effects on the surrounding bone, an alternative with a similar modulus of elasticity to bone is required [1-5]. The target values for the modulus of elasticity are nearly a tenth of the value for the titanium alloys commonly used in biomedical implants. While current implants are most often solid structures, modern research has investigated using lattice structures to reduce the stiffness of implants [2, 6, 7]. To create a structure that approaches the desired modulus of elasticity, a density gradient is usually applied to the final structure [8]. Ultimately, the core of the structure is denser than the outer structure [8, 9].

Ultimately, the core of the structure is denser than the outer structure [8, 9]. This gradient change is discussed by Xu-bin et al [10]. Bio-medical implant data is available in Ref. [11] which is widely-used data source for design of lattices. Shrestha et al. [12] studied the behavior of lattices under tensile loading; however, determination of behavior in compression is crucial for bio-medical implant applications.The work that follows is the preliminary process examining the stress effects of changing volume fraction. This is critical for choosing the most promising lattices as well as determining2196which parameter is most consequential to change regarding the stress profiles of the lattices. Throughout the design and analysis process, a library of the seven lattices was compiled. This library serves as a reference for the research and contains mechanical properties of each lattice’s varied volume fraction.

This library serves as a reference for the research and contains mechanical properties of each lattice’s varied volume fraction. This will assist in deciding which lattice structures exhibit the preferred mechanical properties. Once the data is acquired, it will be used in conjunction with other biomechanical data, to assist in the creation of a patient specific implement that reduces the surgery’s complications as well as improving life after the implant.Architectural Design DevelopmentTo obtain adequate research data, the authors selected seven different unit cells to analyze (see Table 1). All strut-based unit cells analyzed were variants of face centered cubic (FCC) and body centered cubic (BCC) unit cells. They were chosen because of applicability in biomedical implants and manufacturability by additive manufacturing processes.Varying volume fraction allows for a variation in porosity and relative density of the unit cell structure.

Increasing porosity promotes decreased weight and stiffness of the structure. This is desirable because the implant will more closely mimic the stiffness of the bone that it will be integrated with, reducing the effect of stress shielding. Additionally, a lattice structure promotes a higher degree of osseointegration between bone and implant than a solid material because it contains space in which bone tissue can grow.SolidWorks is the software used for modeling and simulation. The process of computer aided design (CAD) file generation for lattice samples started by creating a unit cube, with equal dimensions on the x, y, and z axes and centered at the origin. Construction planes were then created at appropriate orientations from which struts could be extruded to form the unit cell. All dimensional parameters were linked together, lengths based off cubic width and diameters made equal to the first strut extruded.

All dimensional parameters were linked together, lengths based off cubic width and diameters made equal to the first strut extruded. This linkage allows the user to modify parameters of a completed lattice structure by changing just two dimensions. From this point, construction direction lines in x, y, and z axes were created. The linear pattern and mirror features were used to produce a 6x6x6 unit cell lattice structure in those directions. Table 2 illustrates the different lattice structures, and Table 3 shows the varying dimensions of a BCCZ lattice structure. All lattices were initially tested at strut diameters of 0.4, 0.8, 1, 1.5, and 2 mm.Table 1: Unit CellsFCCFCCZF2CCF2CCZBCCBCCZF2BBC2197Table 2: Lattice samplesFCCFCCZF2CCF2CCZBCCBCCZF2BCCIn this step of the study, stainless steel 316L SS was selected for all FEA analysis for the sake of cost effectiveness.

Applicability of this type of bio-compatible metal for this purpose will be discussed later but was selected based on the work done by Chen et al. [13]. Associated material properties needed for the simulation was available by default in SolidWorks.Computing Issues/Mesh AdequacySince this study was heavily based on finite element analysis, ensuring the accuracy of the simulation results is pivotal. An analysis of meshes of different elements sizes was performed to converge to an appropriate mesh. Due to time constraints, an optimally coarse mesh had to be used to save simulation time. The complexity of these 6x6x6 lattice structures caused meshing issues as well.In order to prove the accuracy of the mesh that was selected, a beam structure fixed at both ends was simulated. It was loaded with a single point load at the center and the max stresses and displacements were compared to traditional hand calculations of a beam loaded under the same conditions.

It was loaded with a single point load at the center and the max stresses and displacements were compared to traditional hand calculations of a beam loaded under the same conditions. Mechanics of materials equations (Eqs. (1) and (2)) were used to calculate stress and displacement of this type.Maximum moment: (cid:2026) (cid:3404)(cid:3014)(cid:3052)(cid:3010)(1)Maximum displacement: (cid:1986) (cid:3404)(cid:3017)(cid:3039)(cid:3119) (cid:2869)(cid:2877)(cid:2870)(cid:3006)(cid:3010)(2)The analyzed model is shown in Figure 1(a). The analysis was based on a beam of 10 mm in length and 1 mm in diameter to closely simulate a lattice strut. Identical simulations were run on a variety of meshes until there was a convergence in the stress value. As shown in Figure 1(b), even a mesh with twenty times more elements displayed similar results to the coarser meshes.

As shown in Figure 1(b), even a mesh with twenty times more elements displayed similar results to the coarser meshes. The values still have a slight variation of approximately 5% with respect to the hand calculations2198because the nodes of the mesh were not in the exact same location each time the comparison values were taken [14].Stress Value vs. Number of Elements145000000135000000f l 125000000..f;; 115000000105000000950000002000"""6000Number of Elements800010000(a)(b)Fig. 1. Mesh adequacy evaluation; (a) 3D model of beam, (b) numerical convergence of FE solutionParameter Dependency StudyVolume fraction defined by relative density of the lattice structures to a solid cube of wrought stainless steel is examined as the main parameter.

Various lattice 6*6*6 samples with different sizes but identical volume fractions were constructed for each unit cell (see Table 3 for an illustration of proportional differences and Table 4 for lattice properties). The provided data confirms that the given unit cell lengths and strut diameters produce equal mass and volume fractions. Further FEA was performed to confirm that volume fraction is the only parameter that influences Elastic Modulus.Table 3: BCCZ Lattice Variations.4 mm.8 mm1 mm1.5 mm2 mmThe design studies were created by fixing a central node on the bottom of the lattice, then applying a roller/slider constraint on the rest of the bottom nodes. Roller/slider constraints were used instead of fixed geometry to avoid unnecessary and inaccurate moments being applied to bottom nodes.

Roller/slider constraints were used instead of fixed geometry to avoid unnecessary and inaccurate moments being applied to bottom nodes. The top nodes were subjected to homogeneous prescribed displacement, which was increased incrementally in consecutive studies until the resulting stress reached yield for the material. The same process was repeated across all unit cells and all volume fractions.

The same process was repeated across all unit cells and all volume fractions. Properties219912000were then compared to those of wrought material and effective values for the lattices were extrapolated.Table 4: Mass and Volume Fraction CalculationsUnit Cell Length (mm)Strut Diameter (mm)Cube Mass (g)Lattice Mass (g)Mass FractionCube Volume (mm3)Lattice Volume (mm3)Volume Fraction51216.7335.9 0.16564270004472.06 0.165631102 1733.83287.19 0.1656321600035777.87 0.165638153 5851.68969.27 0.16563729000 120751.19 0.165639204 13870.62297.52 0.16563 1728000 286223.88 0.165638The visual displacement results for this process can be found in Figure 3, while Figure 4 shows the Von Mises stress contour. As expected, the boundary conditions imposed on the models resulted in homogeneous distributions of displacement, and symmetric distributions of stress.

As expected, the boundary conditions imposed on the models resulted in homogeneous distributions of displacement, and symmetric distributions of stress. It is important to note that maximum stress locations, colored red on the stress map, are found where the geometry changes abruptly due to stress concentration at strut intersection. These stress concentrations, or hotspots, are to be expected. The few outliers in the data collected generally were caused by hotspots behaving asymmetrically. Fortunately, these outliers can generally be ignored due to the ability to correlate data with lattices using different volume fraction parameters. In further FEA studies, these hotspots can be mitigated by changing geometry such that strut intersections occur more smoothly.

In further FEA studies, these hotspots can be mitigated by changing geometry such that strut intersections occur more smoothly. Table 5 summarizes the result for parameter dependency study.The data in Table 5 demonstrates that the effective elastic modulus remains constant, within ± 0.84%, when unit cell length and strut diameter parameters are changed while keeping volume fraction equal.

This further proves that volume fraction is the only parameter that effects lattice strength.➔ Yitk1Mf'Ongtl<l.700o•08Figure 1: F2CCZ Displacement StudyFigure 2: F2CCZ Von Mises Study2200Table 5: Effective Elastic Modulus calculationUnit Cell Length (mm)Strut Diameter (mm)Area (mm2)Force (N)Effective Stress (MPa)ΔL (mm*E -03)L (mm)Strain (*E-5)Effective Elastic Modulus (N/m2*E+10)5190017281.925.003016.71.151023600 3485.396.85.00608.331.161538100 5211.264.35.00905.561.1620414400 6939.548.25.001204.171.16Max:1.16E+10Min:1.15E+10%Diff:8.48E-01Based on the procedure discussed in next section, stiffness and yield strength of lattice samples were determined.

Specific strength was also calculated and included in addition to the previous values.Parameter Dependency of Mechanical PropertiesThrough the described finite element method, the effect of volume fraction on the mechanical properties of all seven unit cell lattice variations was investigated. The properties analyzed were effective elastic modulus, effective yield strength, and specific strength (strength- to-weight ratio). After performing many iterations of the described method applying different loads the force versus displacement curve (an example in Fig. 5) is obtained. Then apparent stress and strain are calculated and Elastic Modulus is determined as the slope of stress-strain line. Yield stress is also considered as the apparent stress at onset of exceeding maximum Misses stress of the structure from yield strength of 316L SS. If the structure meets this condition, it begins to deform permanently.

If the structure meets this condition, it begins to deform permanently. As an example, the force versus displacement diagram for the BCC lattice structure with a 1.06 mm diameter is presented in Figure 5. Figures 6 to 8 summarize the data generated as the library of unit cells which contains variation of stiffness and strength with volume fraction for each type of cells.By generating this data and understanding these relationships, the bio-implant designers can optimize specific mechanical properties based on the application of the product being designed. If a design requirement is weight, stiffness, strength, or some combination of all, they have the capability with this study and method to vary volume fraction (unit cell geometry and type) in a particular part and achieve desired properties in all locations of the product.22014500I400035003000.....

V,--+- FCCZ+- FCC--+- F2CCF2CC2--+- F2BCC--+- BCC2+- sec2.00E+031.00E+030.OOE+OO00 .10 .20 .30 .40 .50 .60 .70 .80.9Volume FractionFigure 8: Specific Strength ComparisonPractical Application of Lattice LibraryReferring to the data provided by Zhang, et al. [5] for human bones, it was attempted to find proper lattices as the replacement for bones with the application in bio-implants. The proper volume fractions for each unit cell are defined by matching stiffness from Ref. [5] with associated ones in2203our library of lattices (Fig. 6). The corresponding unit cell volume fractions are shown in Table 6.

Considering the higher strength of lattices in tension (3-5 time larger), they are still significantly weak for this application (see Fig. 7). The proposed solution is changing material from 316L SS to Ti6Al4V which is widely-used standard material for biomedical implants, referring to Ref. [15].ExperimentComputer aided manufacturing (CAM) files were generated at National Center for Additive Manufacturing Excellence (NCAME) for all seven lattices using Materialise Magics software and then sent to EOS 290 metal 3D printer to fabricate samples for testing. Due to some manufacturability issues Ti64 is considered for printing the samples and those simulations used for evaluation of FE models are to be repeated with appropriate material data, consequently. Figure 9 shows some of the fabricated lattice compression samples and Figure 10 indicates the compression test setup.Fig. 9. Lattice compression samples designed and manufactured in NCAME2204Fig. 10.

9. Lattice compression samples designed and manufactured in NCAME2204Fig. 10. Setup for compression testTables 7 and 8 summarize the result obtained from two tests for Elastic Modulus and maximum force. Maximum force is considered as the force which creates plastic deformation in structure.Table 7: Measured Elastic Modulus in experimentsBCCElastic Modulus Test Results BCCZFCCFCCZF2CCZTest 1 (GPa)0.40.9060.71I.II1.43Test 2 (GPa)0.4NIA0.711.161.58% Difference0.00%NIA0.71 %4.50%10.49%Table 8: Measured maximum forces in experimentsMaximum Force Test ResultsBCCBCCZFCCFCCZF2CCZTest 1 (kN}19.7239.930.6561.3368.8Test 2 (kN)20.66NIA31.5 161.3167.99% Difference4.55%NIA2.73%0.03%1.19%As it can be seen in Tables 7 and 8, the results obtained for two tests are in a good agreement and repeatability of experiment is guaranteed.

These data will be used for evaluation purposes in the next step of this comprehensive study.2205F2CC1.061.169.43%F2CC46.9247.962. 17%ConclusionThe research-based approach of studying these lattice structures lead the authors to investigate the effect of volume fraction on elastic modulus, yield strength, and specific strength. This was done because the volume fraction study that was performed indicated that the main factor affecting these mechanical properties is volume fraction and not individual dimensions. These issues caused limitations to the mesh capabilities for the simulations, so an in-depth mesh analysis was performed to ensure that coarser meshes presented similar results to finer meshes.At its current state, the project has generalized a design optimization method by utilizing the relationships between mechanical properties and the geometries of a particular lattice structure.

The process provides designers with the ability to vary structures with volume fraction to achieve desired properties in specific locations of a product. In practical application, the stiffness can be matched and proper lattices can be determined as the replacement of different human bones in bio- medical implants. However, there is an issue with matching strength. Selected lattice structures made from 316L SS are significantly weaker than associated bones. Therefore, Ti6Al4V with higher strength and lower density would be used to fabricate lattices in further steps of this study.References1. Balla, Vamsi Krishna, et al. "Porous Tantalum Structures for Bone Implants: Fabrication, Mechanical and In Vitro Biological Properties." Acta Bomaterialia, vol. 6, no. 8, Aug. 2010, pp. 3349-59, doi:10.1016/j.actbio.2010.01.046. Accessed 26 Feb. 2019.2. He, Yuhao, et al.

mputational Design, Freeform Fabrication and Testing of Nylon-6 Tissue Engineering ScaffoldsSuman Das†, Scott J. Hollister‡, Colleen Flanagan‡, Adebisi Adewunmi†, Karlin Bark†, Cindy Chen†, Krishnan Ramaswamy†, Daniel Rose†, Erwin Widjaja††Mechanical Engineering Department ‡Biomedical Engineering Department University of Michigan Ann ArborAbstractApproximately 100,000 people in the US alone suffer from TMJ disease to the extent that surgical reconstruction is needed. In order to reconstruct a whole joint such as the TMJ, advanced, novel fabrication methods are needed to build complex, three- dimensional scaffolds incorporating multiple, functionally graded biomaterials and porosity that will enable the simultaneous growth of multiple tissues including blood vessels. The aim of this research is to develop, demonstrate and characterize techniques for fabricating such scaffolds by combining solid freeform fabrication and computational design methods.

The aim of this research is to develop, demonstrate and characterize techniques for fabricating such scaffolds by combining solid freeform fabrication and computational design methods. When fully developed, such techniques are expected to enable the fabrication of tissue engineering scaffolds endowed with functionally graded material composition and porosity exhibiting sharp or smooth gradients. As a first step towards realizing this goal, we have designed and fabricated scaffolds with various architectures in Nylon-6, a biocompatible polymer using selective laser sintering. Results of bio- compatibility, mechanical testing and implantation are presented.INTRODUCTION Tissue engineering scaffolds aim to serve three purposes: 1) They form regenerate tissue shape. 2) They temporarily fulfill the physical function (mechanical, chemical and/or electrical) of the native tissue.

2) They temporarily fulfill the physical function (mechanical, chemical and/or electrical) of the native tissue. 3) Their material composition and porous microstructure enhance tissue regeneration and vascularization.Polymers based on polylactic acid (PLA), polyglycolic acid (PGA) and their co- polymers (PLGA) are the most widely investigated biocompatible, biosorbable polymers for bioimplants [1,2] and controlled release drug delivery devices [3]. Composites of these polymers with hydroxyapatite (HA) are being investigated as bone fixation devices in orthopedic, craniofacial, maxillofacial and reconstructive surgeries. PMMA, Nylon-6 [4], Polycaprolactone [5], and bioactive glass [6] are being investigated as well.Current methods of scaffold/bioimplant fabrication include particulate leaching, solvent casting, forging, injection molding and cold or hot pressing. These methods provide limited control over scaffold architecture and pore geometry.

These methods provide limited control over scaffold architecture and pore geometry. More recently, SFF methods have been employed for fabricating bioimplants and tissue engineering scaffolds. Chu et al. [7] have fabricated designed porous bio-implants by ceramic stereolithography. Lewis et al. [8] are exploring fabrication of mesoscale 3-D periodic structures for biomaterial scaffolds with locally tailored composition using the robocasting process. Giordano et al [9] have used 3D printing to fabricate scaffolds out of PLA, PGA and PLGA. Lee et al [10] have used selective laser sintering (SLS) to construct artificial bone implants from hydroxyapatite and calcium phosphate ceramics. Taboas et al are using thermojet printing to fabricate molds for casting PLA, PLGA and9hydroxyapatite scaffolds.

Taboas et al are using thermojet printing to fabricate molds for casting PLA, PLGA and9hydroxyapatite scaffolds. Zein et al [5] are using fused deposition modeling to construct scaffolds from polycaprolocatone.Several hypotheses propose that 1) the architecture and composition of scaffolds can affect tissue regeneration. 2) cells behave differently in a composite scaffold with graded composition and porosity when compared to a monolithic scaffold with uniform porosity. 3) scaffolds with graded material and porosity interfaces are capable of generating multiple different tissues simultaneously and are important for biological systems with tissue interfaces. However, these hypotheses have not been tested to date due to lack of heterogeneous that scaffold fabrication methods.

However, these hypotheses have not been tested to date due to lack of heterogeneous that scaffold fabrication methods. Additional hypotheses state architectures must be designed to have optimal pore size and structure to promote specific tissue regeneration while matching mechanical properties of native tissue as closely as possible. In principle, SFF methods are capable of constructing three- dimensional scaffolds with complex architectures incorporating multiple, functionally graded bio-materials and porosity and are therefore ideally suited to test these hypotheses. The overall goal of this research is to develop homogenization theory based computational design and laser sintering methods for constructing such heterogeneous scaffolds using biopolymers and bioceramics. As a first step towards realizing this goal, we have chosen to investigate monolithic Nylon-6 as a scaffold material for tissue engineering applications.OBJECTIVESThe objectives of this research are to: 1.

locally optimize scaffold Develop computational architecture, material composition and mechanical properties yielding three- dimensional digital representations of functionally tailored scaffolds. Develop solid freeform fabrication techniques based on laser sintering that can construct scaffolds using multiple biomaterials possibly with drug/biofactor delivery. Conduct mechanical and biological (in vivo, in vitro) testing, and CT/MRI analysis to evaluate structure and function of both scaffold materials and regenerate tissue.techniquesthat can2.such3.METHODS AND MATERIALS Scaffold Design MethodsScaffold design requirements must be addressed on both macroscopic and microscopic scales. On the microscopic scale (100µm-1mm), the scaffold internal architecture must a) Fulfill temporary tissue function b) Enhance tissue regeneration and vascularization and c) Facilititate nutrient/biofactor delivery.

On the macroscopic scale (> 1mm) , the scaffold's external shape must replicate human anatomy. These two scales must be integrated to produce a single design that can be embodied in a format appropriate for SFF. We have used both periodic cell-based designs and biomimetic designs to construct scaffolds using SLS.10Periodic Cell-Based DesignsIn periodic cell-based designs, a unit cell with specific microstructure is repeated to create an entire scaffold. This technique can be used in combination with topology optimization methods to design microstructures with effective physical properties matching native tissue properties. A scaffold unit-cell microstructure, shown in Figure 1 comprises intersecting porous cylinders and is characterized by three pore diameters.

A scaffold unit-cell microstructure, shown in Figure 1 comprises intersecting porous cylinders and is characterized by three pore diameters. Homogenization theory is used to compute a functional dependence of effective scaffold property Ceff scaf and effective regenerate tissue property Ceff regen tissue on scaffold base property Cscaf, scaffold microstructure Mscaf, regenerate tissue base property Cregen tissue and regenerate tissue microstructure Mregen tissue.

The dependence of effective scaffold property on scaffold base property and scaffold microstructure can be written asCeff scaf=CscafMscaf(d d d 1 2,,3)Ceff regen tissue=Cregen tissueMregen tissue(d d d 1 2,,3)Using the optimization formulation shown below, scaffold pore geometry is designed so that the effective scaffold stiffness and the effective regenerate tissue stiffness match native tissue stiffness as closely as possible while meeting constraints on the base scaffold material stiffness, on scaffold porosity, and on pore diameters.Objective Function:       Constraints: ≤ , d d d 1 2 d d d 1 2 VMin , , d d d 1 2scafC,3,3≥,,3n= → 1∑=i19  −eff regen tissue C ieff native tissue C i eff native tissue C in= → 1∑=i19  −eff scaf C ieff native tissue C i eff native tissue C i  2largest pore size for regenerationresolution of SFF technique  2 +     poreVtotal≥% PorosityCscaf≥CminCscaf≤CmaxFigure 1.

Scaffold unit-cellFigure 2 shows 8mm cubic and 8mm diameter cylindrical periodic scaffolds with 800µm orthogonal channels and 1200µm pillars that were designed using the above optimization methods.11Figure 2. Cubical and cylindrical periodic scaffold architecturesBiomimetic DesignsThe design of biomimetic scaffolds relies on micro-CT, micro-MRI or confocal microscopy data to assemble scaffold architectures. In biomimetic designs, scaffolds mimic natural tissue structure and seek to replicate all aspects of tissue structure and function. They are most difficult to achieve due to complexity of biological tissues, which are composites organized in a hierarchy from nanometers to millimeters. We have used human trabecular bone architecture as the basis for creating biomimetic scaffold microstructures. Figure 3 shows a volumetric rendering of human trabecular bone micro- CT data along with a faceted representation appropriate for use in SFF.Figure 3.

Figure 3 shows a volumetric rendering of human trabecular bone micro- CT data along with a faceted representation appropriate for use in SFF.Figure 3. Trabecular bone microstructure used for biomimetic design. (a) Shaded rendering of image data. (b) Triangular facet data for SFF.Scaffold FabricationNylon-6 was chosen as the material for fabricating scaffolds during this study. This choice was prompted by Risbud and Bhonde’s data [] on the biocompatibility of polyamide 6. Their aim was to develop polyamide 6 membranes blended with gelatin (a natural polyamide) and chondroitin sulfate (a biopolymer) using the phase precipitation method and evaluate its in vitro biocompatibility. A large collection of biocompatibility test data demonstrated that polyamide 6 composite membranes are biocompatible and prospective candidates for tissue engineering.

A large collection of biocompatibility test data demonstrated that polyamide 6 composite membranes are biocompatible and prospective candidates for tissue engineering. Additionally, Nylon has been routinely used in suture materials, for dialysis membranes, burn dressings, and cell culture substrata.12Nylon-6 powder (10-100µm particle size) used in our study was prepared by cryogenic milling of pellets. The powder was processed in an SLS Sinterstation 2000 using 200º C preheat, 7 Watts laser power, 49.5 in/s scan speed and 100 microns layer thickness. Porous specimens of both cylindrical and cubical geometry (Figure 2) were fabricated.Mechanical TestingA modified ASTM D638 Type II test specimen (Figure 4) was designed by incorporating a porous architecture (700µm orthogonal porosity, 800µm walls) in the gauge section of the tensile bar. The tensile bar is 9.8mm thick, 13mm wide, and has a 57mm long gauge section.

The tensile bar is 9.8mm thick, 13mm wide, and has a 57mm long gauge section. Mechanical testing was conducted to determine tensile strength and elastic modulus for porous and non-porous specimensFigure 4. Modified ASTM D638 type II specimenRESULTS AND DISCUSSIONFigure 5 shows an 8 mm cube with 800 micron channels and 1200 micron pillars fabricated in Nylon-6 by SLS. These specimens will be used for conducting uniaxial tests in unconfined compression inside a micro-CT machine. Complete 3D strain fields in the scaffold under testing will be computed by comparing images before and after deformation. These tests will provide effective failure stress and local strain values at failure.Figure 5. 8mm cubic periodic scaffold fabricated in Nylon-6Figure 6a shows an 8mm diameter, 6mm high cylinder with 800µm channels and 1200µm pillars fabricated in Nylon-6 by SLS. This scaffold geometry was designed for surgical implantation and histology assessment.

This scaffold geometry was designed for surgical implantation and histology assessment. Figure 6b shows implantation of cylindrical scaffolds into a pig. These scaffolds were subsequently removed after 6 weeks and assessed for mineralized tissue formation. Figure 7 shows a volumetric rendering of the micro-CT scan conducted on the removed scaffold, showing the growth of mineralized tissue into the pore channels of the scaffold.13Figure 6 (a) 8mm diameter, 6mm high periodic cylindrical scaffold fabricated in Nylon-6 and (b) scaffold implantation into pig.Figure 7. Volumetric rendering of scaffold micro-CT scan.Biomimetic structures derived from CT/MRI data are difficult to fabricate by SFF as resolved tissue structures (10-100µm) are often smaller than resolution of SFF machines. Alternatively, these structures can be scaled up and then fabricated to have optimal pore size for tissue regeneration (typically 300-1200µm) while retaining biomimetic architecture.

Alternatively, these structures can be scaled up and then fabricated to have optimal pore size for tissue regeneration (typically 300-1200µm) while retaining biomimetic architecture. Shown in Figure 8 are the volumetric rendering of human trabecular bone micro-CT data and the corresponding Nylon-6 replica scaled 4X fabricated using SLS. In addition to biological testing of these scaffolds by implantation, we will visualize 3D deformation and failure modes under compression during micro-CT and compare them with failure modes of real bone.Figure 8 (a) volumetric rendering of human trabecular bone micro-CT data and (b) a 4X scaled replica fabricated in Nylon-6 by SLS (scale in mm).14Mechanical PropertiesMechanical properties of porous and non-porous Nylon-6 tensile specimens areshown in Table 1.Specimen Non-porous PorousUltimate Tensile Strength (MPa) 42.7±2.5 13.4±0.7Modulus of Elasticity (MPa) 503±37.2 250.3±10.2Table 1.

Mechanical properties of non-porous and porous Nylon-6 tensile specimensModulus of elasticity (effective stiffness) for cancellous (or trabecular) bone tissue exhibits a wide range of values due to a wide variety of bone architectures. Reported values for measured effective stiffness are in the 70-445 MPa range whereas computed effective stiffnesses are reported to in the 50-2000 MPa. Therefore, the elastic moduli of both porous and non-porous Nylon-6 specimens are consistent with stiffness of trabecular bone. While it is very difficult to measure tensile properties of cortical bone, compressive ultimate stress for cortical bone is reported in the 2.2-7.4 MPa. Therefore, the ultimate strength values in tension for Nylon-6 are larger than reported values for trabecular bone effective ultimate compressive strength. Reported values for cortical bone stiffness are in the 12-20 GPa range whereas ultimate tensile strength is reported in the 50-150 MPa range.

Reported values for cortical bone stiffness are in the 12-20 GPa range whereas ultimate tensile strength is reported in the 50-150 MPa range. Therefore, properties for both non-porous and porous Nylon-6 are far below those of cortical bone.A homogenization finite element analysis of Nylon-6 porous architecture with 700µm porosity and 800µm wall thickness was conducted to computationally validate the experimentally measured effective properties of porous specimens. Experimentally measured stiffness of 503 MPa for solid Nylon-6 was used as the base material stiffness in the computations. The material was assumed isotropic with 0.3 Poisson’s ratio. A voxel finite element model with 27,000 solid 8-node hexahedral elements (30x30x30) yielded an effective stiffness of 221.6 MPa while a more refined model with 125,000 elements (50x50x50) using the same analysis yielded an effective stiffness of 224.1 MPa.

Therefore, the homogenization finite element analysis results compare very well with experimentally measured values in the 238-257 MPa range with mean value of 250.3 MPa.CONCLUSIONS An approach combining computational design, freeform fabrication and testing of tissue engineering scaffolds was demonstrated. Scaffolds were fabricated in Nylon-6 using SLS. Implantation and subsequent histology of scaffolds show presence of regenerate mineralized tissue. Mechanical properties of scaffold architecture were measured using a modified ASTM D638 test. The measured elastic moduli are consistent with those of trabecular bone and are confirmed by finite element analysis. Although not biosorbable, such Nylon-6 scaffolds are biocompatible and could serve as drug/cell delivery devices as well as models for future work on PLA, PGA, PLGA. This work sets the stage for construction of functionally tailored tissue scaffolds in a single step via SLS of multiple materials.

This work sets the stage for construction of functionally tailored tissue scaffolds in a single step via SLS of multiple materials. We are exploring graded architectures with multiple biopolymers and their composites with calcium phosphate ceramics. A multiple material deposition system is being developed to make this fabrication approach successful.15REFERENCES 1. A. Rothen-Weinhold et al., Injection-molding versus extrusion as manufacturing technique for the preparation of biodegradable implants, European Journal of Pharmaceutics and Biopharmaceutics, Vol. 48, 1999, pp. 113-121.2. F. Stancari et al., Use of PLA-PGA(Copolymerized Polylactic/Polyglycolid acids) as a bone filler: Clinical Experience and Histologic Study of a Case, Quintessenz (Germany) 51, 1, 47-52, 2000.3. Scott D. Putney, Advances in Protein Drug Delivery, Pharmaceutical News, Vol. 6, No. 2, 1999.4. Makarand V. Risbud and Ramesh R.

MPUTATIONAL MODELING AND EXPERIMENTAL VALIDATION OF MELTING AND SOLIDIFICATION IN EQUIAXED SUPERALLOYS PROCESSED THROUGH SCANNING LASER EPITAXYAmrita Basak*, Ranadip Acharya*, Rohan Bansal*, and Suman Das*George W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332AbstractThis paper focuses on simulation-based optimization of the Scanning Laser Epitaxy (SLE) process applied to gas turbine hot-section components made of nickel-base superalloys. SLE creates equiaxed, directionally-solidified and single-crystal microstructures from superalloy powders melted onto like-chemistry substrates using a fast scanning, high power laser beam. In this paper, a transient coupled flow-thermal approach is implemented to accurately simulate the melting and solidification process in SLE. The laser movement is modeled as a Gaussian moving heat source, and the thermophysical properties of the alloys are adjusted based on the thermal field.

The laser movement is modeled as a Gaussian moving heat source, and the thermophysical properties of the alloys are adjusted based on the thermal field. Simulations for different superalloys such as IN100, René 80 and MAR-M247 are performed and the instantaneous melt pool characteristics are recorded. Comparisons of the simulations with experimental results show reasonably good agreement for the melt depth. Feedback control is implemented, and demonstrated to produce superior quality SLE deposits. This work is sponsored by the ONR through grants N00014-11-1-0670 and N00014-14-1-0658.IntroductionThe design life of commercial gas-turbine engine hot-section components is typically shorter than their cold section counterparts, primarily because of their significantly elevated operating temperatures.

The ever increasing demand for operating modern state-of-the-art engines at higher temperatures has put forward a challenging task for engineers to design suitable materials for these components. Material issues are of paramount significance for these components. Currently nickel-base superalloys are used to manufacture these parts as they offer higher yield strength with increasing operating temperatures primarily due to the presence of the secondary γ’ phases that form an antiphase boundary and lock the associated dislocation movements. The degradation of the hot-section components are attributed to creep, thermo- mechanical fatigue, high temperature oxidation, and thermal corrosion. Currently, the lack of suitable commercial repair technologies results in high maintenance costs for these components that can reach as much as hundreds of thousands to even millions of dollars per engine.

Hence, there is a great commercial interest in additive manufacturing processes capable of restoring the parent metallurgical microstructure and geometry at damage locations for these parts. This paper focuses on simulation-based process development for scanning laser epitaxy (SLE); a laser based additive manufacturing process that has demonstrated the capability of repairing these high-value parts.SLE is a metal powder bed fusion-based additive manufacturing process that creates equiaxed (EQ), directionally solidified (DS) and single-crystal (SX) structures in nickel superalloys through controlled laser melting of alloy powders onto like chemistry substrates. InSLE, a galvanometer controlled high power laser beam is directed onto the powder bed. Processing parameters such as laser power, preheat and scan speed are varied in SLE to control and optimize the deposit characteristics.

Processing parameters such as laser power, preheat and scan speed are varied in SLE to control and optimize the deposit characteristics. An important criterion for a successful SLE deposit is the determination of energy density that ensures a good metallurgical bond between the substrate and the deposit along the entire length of the substrate. The metallurgical bond starts at one edge of the substrate where the melt pool is initiated. The melt pool travels across the length of the substrate fusing the powder to the substrate. Under the proper operating conditions, the solidified microstructure has been demonstrated to nucleate from and follow the morphology of the underlying substrate.Computational modeling and simulation of the SLE process can provide critical insights into the physics of the process and help in determining the desired operating parameters for different superalloy powders, thereby reducing the number of experimental runs.

Several reports on the simulation of laser-based additive manufacturing approaches can be found in literature. Both Finite Element Methods (FEM) and Finite Volume Methods (FVM) are used for the thermal modeling of the system [1, 2]. Gäumann implemented a modified Rosenthal solution [3] and obtained the liquid isotherm and a circular melt pool [4]. Christensen et al. has suggested the formation of an elliptic melt pool along the direction of the beam’s movement [5]. Goldak et al. assumed a double-ellipsoid weld pool shape that consists of front and rear ellipsoid quadrants [6]. A coupled thermal-flow model allows incorporation of fluid convection inside the melt pool. To model the flow pattern prevailing inside the melt pool, surface tension driven flow known as Marangoni convection (or Bernard-Marangoni convection) is typically included in the flow- thermal model [7]. Inclusion of Marangoni convection extends the melt pool laterally and makes it shallower [8, 9]. Acharya et al.

Inclusion of Marangoni convection extends the melt pool laterally and makes it shallower [8, 9]. Acharya et al. [10] developed a three-dimensional CFD model and obtained good agreement with experimental data for the melt depth in single-crystal alloy CMSX-4 processed through SLE. Acharya et al. implemented a pseudo-2D approach to model the melt pool characteristics for IN100 and obtained reasonably good agreement for the melt depth and the average melt pool temperature [11].In this work, Acharya’s pseudo-2D CFD model [11] has been extended to simulate and compare the melt pool characteristics for different equiaxed superalloys. The movement of the laser beam along the x and y directions is modeled as a line source scanning in the x direction. The laser power profile in the y-direction is modeled from the transient averaged data for a single scan and approximated with a 10th degree polynomial [11].

The laser power profile in the y-direction is modeled from the transient averaged data for a single scan and approximated with a 10th degree polynomial [11]. This polynomial is found to be flat in the middle region with some reduction of intensity near the edges of the sample. This characteristic of the power profile justifies the simplification of the computationally expensive three-dimensional CFD model to a tractable two-dimensional simulation. The width of the coupon is used to calculate the residence time of the line heat source at any given position, and then the heat source is advanced in the forward direction by the scan spacing. The complete scan pattern is formulated using ANSYS CFX user-defined functions. The heat conduction equation is used to model the transfer of heat to the base plate. Analysis of the flow-field modeling reveals the formation of rotational vortices due to the moving heat source and the Boussinesq convection.

Analysis of the flow-field modeling reveals the formation of rotational vortices due to the moving heat source and the Boussinesq convection. Reasonable agreement between the simulation and the experimental results is achieved for the melt depth for different superalloy powders. The effect of SLE operating parameters on the melt depth has been analyzed and presented in this work.SLE Experimental ProcedureSLE processing of the nickel-base superalloys such as IN100, René 80 and MARM-247 is investigated in this work. The René 80 and MARM-247 powders are produced by Praxair Surface Technologies using an atomization process. The IN100 powder is produced by HMI Metal Powders using Argon gas atomization. Table 1 summarizes the chemical composition of these superalloy powders.Table 1. Chemical composition of various superalloys (wt.

Table 1 summarizes the chemical composition of these superalloy powders.Table 1. Chemical composition of various superalloys (wt. %)Cr Co Mo Re W Al Ti Ta Hf - - 10.0 15.0 3.0 4.0 14.0 10.0 4.0 - 10.0 6.0 1.0 3.0 1.0 -IN100 René 80- -4.7 5.5 3.0 5.0-MARM-247 8.0 10.0Zr Ni 0.014 0.18 0.06 Bal Bal BalBC---The SLE process is conducted on EQ cast rectangular coupons. Coupon dimensions for IN100 and René 80 are 35.56mm X 6.86mm X 2.54mm while for MAR-M247 the coupons are slightly smaller in size having dimensions of 31.50mm X 5.84mm X 2.54 mm. Each substrate is placed into a 35.56mm x 6.86mm (for IN100 and René 80) and 31.50mm X 5.84mm (for MAR- M247) recess cut into an Inconel 625 base plate. A powder bed is placed atop the substrate using rectangular wells cut into an Aluminum mask plate. Once the samples are prepared, they are placed into an atmospheric process chamber that is purged with high purity (99.999%) Argon.

Once the samples are prepared, they are placed into an atmospheric process chamber that is purged with high purity (99.999%) Argon. A 1kW Ytterbium fiber laser (IPG Photonics, Model: YLS-1000) is used along with a Cambridge Technologies galvanometer scanner to scan the beam on the powder bed. The beam is focused to a 40 µm Gaussian diameter. A raster scan pattern across the width of the sample generates a melt pool that linearly propagates along the length of the substrate. The details of the experimental setup can be found elsewhere [11], and is skipped here for brevity.SLE Deposit CharacterizationSLE-processed samples are sectioned along the length and width for metallography. A Buehler automated saw is used to cut the samples. A Leica DM6000 optical microscope is thereafter used to take the images.

A Buehler automated saw is used to cut the samples. A Leica DM6000 optical microscope is thereafter used to take the images. Figure 1 (a) and 1 (b) demonstrate lengthwise cross-sectional view of crack-free IN100 and René 80 samples as processed by SLE showing a full metallurgical bond along the entire length of the sample.Figure 1. Representative lengthwise section of the first half of a (a) IN100 and (b) René 80 samples with the starting edge on the left side of the image.Computational Flow-Thermal Modeling of SLEA transient, multi-domain model with temperature-dependent property values is used to simulate the flow-thermal behavior of the SLE and to predict the melt pool characteristics. Temperature-dependent thermo-physical property data for IN100, René 80 and MAR-M247 is modeled using tabular data [11, 12]. The pseudo-2D model includes two separate domains – (1) the deposit domain consisting of the substrate and the powder and (2) the Inconel baseplate domain.

The pseudo-2D model includes two separate domains – (1) the deposit domain consisting of the substrate and the powder and (2) the Inconel baseplate domain. Figure 2 shows the geometry and the mesh for the 2D configuration. The deposit domain has a finer mesh to accurately capture the 20 micron radius laser heat source. The total node count of the mesh is close to 137k and the skewness is below 0.2. The simulation is performed in ANSYS CFX solver.Figure 2. ANSYS CFX simulation model showing different domains – Deposit domain and Baseplate domain.The solid and liquid component enthalpy equations, expressed in terms of volumefraction, r, are:∂hr ( ρ SS S t ∂)∇+( ρhUr SS SS)∇=(kr SS∇T S)+I∂hr ( ρ LLL t ∂)∇+( ρhUr LL LL)∇=(kr LL∇TL)−IHere ρ is the density, h is the enthalpy (including the latent heat), T is the temperature, U is the velocity and I is the interphase heat transfer.

The subscripts S and L refer to the solid and(1a)(1b)liquid components respectively. Equations (1a) and (1b) can also be expressed more compactly as:∂h ( ρ t ∂)∇+( ρhU)Tk •∇−∇•∇=()( ρhUUY −()(SSL−hS))+SHere ρ is mixture density, h is the mixture enthalpy, k is mixture conductivity, YS is the mass fraction of solid [13]. The final term of Equation (2) represents an additional source due to the difference in velocity between the solid and liquid components. This term has a finite value in the mushy region of the flow (0< YS, YL <1). In the regions of pure liquid or pure solid, the mixture enthalpy equation reduces to the appropriate component enthalpy equation. The volume fractions of liquid and solid component are incorporated using tabular data representing solid and liquid fraction of the superalloys in the melting range.

The volume fractions of liquid and solid component are incorporated using tabular data representing solid and liquid fraction of the superalloys in the melting range. The Boussinesq approximation is implemented to model the effect of variable liquid density.A Darcy like source term is incorporated to model the resistance to flow in the mushyregion. The is modelled To prevent division by zero, the term is bounded by a user-specified maximum, C.permeabilityusing Kozeny-CarmanequationMS=min µ  L C ,  K01(−Lr 3 Lr)2  SUU ( −  )In the turbulence equations, sink terms similar to the Darcy term in the momentumequation are applied to damp the turbulence in solid regions [13].kS−=µ LK01(−Lf 3 Lf)2k,S ε−=µ LK01(−Lf 3 Lf)2εFor the baseplate domain, the conduction heat equation is solved.( H ∂ ρ t ∂)) ( TK.

+∇∇=S0Here ρ is the density, H is the enthalpy, K is the thermal conductivity and S0 is thesource/sink term.The conservation of thermal energy is achieved at all interfaces. The convection heat loss is modeled as a surface sink term using Newton’s law of cooling. Radiation loss is modeled using a sink term. The transient model is solved using a time-step value approximately equal to 0.002 s. The inclusion of convection, phase change model and Darcy sink term requires a smaller time-step value for solver stability. The convergence is assured once the RMS value of residual is reduced below 10e-4. Once solved, the result is post-processed in CFD-post to analyze the melt pool nature, the melt depth profile and the temperature distribution.(2)[14].(3)(4)(5)Results and DiscussionComputational simulation of the SLE process sheds light on the flow-thermal field generated by the moving heat source as it traverses across the substrate.

The flow field affects the melt pool formation and the temperature distribution within the melt pool. Figures 3(a) and 3(b) show the velocity and temperature distribution inside the melt pool for IN100 respectively. The melt pool is identified in the regions with liquid IN100 mass fraction of 0.99 or more. This calculation is validated by the fact that inside the melt pool, the temperature is greater than 1639K, the liquidus temperature of IN100.(a)(b)Figure 3. (a) Contour of velocity, (b) Contour of temperature inside the melt pool for IN100.Figures 4(a) and 4(b) show the zoomed view of the temperature contour near the laser source with and without velocity streamlines superimposed. The melt pool consists of two counter rotating vortices with their center located approximately at the location of the laser heat source. The flow field draws loose powder at the top and feeds the melt pool as it expands.(a)(b)Figure 4.

Several CFD simulations are performed with varying SLE processing parameters and the melt depth profile is obtained for IN100. The parameter range that is explored in this study includes repeat scans of 75-125, scan speeds of 300-500 mm/sec, laser powers of 300-500 W. All samples are run with a scan spacing of 25.4 µm to account for the smaller beam diameter of the fiber laser. The effect of these parameters on the CFD-simulated average melt depth is analyzed and presented in Fig. 6. The blue dotted lines represent the 95% confidence intervals for the experimental data [11], black circles represent the CFD simulated data points, and the black line denotes the best-fit line generated from the CFD simulated data points. The nominal SLE run is performed with number of repeat scans of 100, scan speed of 400 mm/s and laser power of 400 W. The experimental and the CFD data points for the melt pool depth are normalized by the corresponding nominal SLE run.

The experimental and the CFD data points for the melt pool depth are normalized by the corresponding nominal SLE run. The prediction shows excellent agreement with the experimental data across the entire range of the experimentaldata points. It can be seen from Fig. 6 (b) that laser power impacts the melt pool depth significantly. With increasing laser power, the melt pool extends laterally and transversely, thereby increasing the melt depth. Increasing the scan speed adversely affects the melt depth [Fig. 6 (a)]. This is primarily due to the fact that with increasing the scan speed, the laser heat source spends less time per unit surface area in processing the substrate and powder; and therefore the melt depth decreases. Increasing the number of repeat scans increases the melt pool depth at the leading edge of the substrate, and therefore increases the average melt depth [Fig.

Figure 7 shows the relation between the melt pool area and the melt pool depth for a complete scan of an IN100 sample. Since the dominant thermal gradient is along the vertical direction, an increase in the melt pool area is expected to increase the melt depth. The linear relation in Fig. 7 reveals that if the melt pool area is kept constant at a desired value, a uniform melt depth can possibly be achieved. The plot also shows the relation between the average melt pool temperature and the melt pool depth for the same sample. Again, the linear relation reveals that if the average melt pool temperature is kept constant at a desired value, a uniform melt depth can possibly be achieved.The objective of the feedback control trials is aimed at determining the effect of maintaining a constant average melt pool temperature on the deposit characteristics.

Data from a thermal imaging camera is used to compute the average melt pool temperature and an adaptive controller is programmed to adjust the laser power accordingly [15]. The scan speed and the number of repeat scans are not changed during the run. The adaptive control trials are able to appropriately control the laser power and thus the average melt pool temperature during SLE processing and produce uniform deposits. As shown in Figs. 8 (a) and 8 (b), both melt depth uniformity and deposit height uniformity is achieved due to control of the average melt pool temperature in Fig. 8(b) compared to a sample processed under open loop control in Fig. (a).500Figure 7. Variations of the melt pool area and average melt pool temperature with the meltdepth for IN100.Figure 8.

8(b) compared to a sample processed under open loop control in Fig. (a).500Figure 7. Variations of the melt pool area and average melt pool temperature with the meltdepth for IN100.Figure 8. Transverse cross section of a representative (a) open-loop specimen and (b) closed-loop specimen for IN100.ConclusionsIn the current work, a two dimensional CFD-based modeling scheme is implemented to simulate the SLE process using a single platform for different equiaxed superalloys. The simulation work is compared with the experimental melt pool depth profile and reasonably good agreement is achieved for IN100, René 80 and MAR-M247. Laser power, scan speed and number of repeat scans are varied for each experimental runs and their effects on the SLE deposit have been analyzed. Excellent agreement is obtained between simulation and experimental results for the melt depth for wide range of SLE processing parameters.

Excellent agreement is obtained between simulation and experimental results for the melt depth for wide range of SLE processing parameters. Feedback control is implemented to accurately control the average melt pool temperature, showing excellent results in maintaining a uniform deposit as well as a uniform melt depth. A future objective is to develop a simulation map that will allow accurate prediction of the microstructure development as a function of operating parameters and melt pool conditions in superalloys processed through SLE.AcknowledgmentsThis work is sponsored by the Office of Naval Research through grants N00014-11-1- 0670 and N00014-14-1-0658. The first author would like to thank Shashank Holenarasipura Raghu for his assistance with metallography.References1. Anderson TD, DuPont JN, DebRoy T. Origin of Stray Grain Formation in Single-Crystal Superalloy Weld Pools from Heat Transfer and Fluid Flow Modeling. Acta Materialia. 2010; 58(4):1441-54.2.

For any given position of the beam, the geometrical parameters of and the temperature gradient in the melt pool are used to estimate the resulting solidification microstructure. Detailed study of the flow field also revealed formation of rotational vortices in the melt pool. Microstructural predictions are shown to be in good agreement with experimental metallography. This work is sponsored by the Office of Naval Research through grant N00014-11-1-0670.IntroductionIn order to improve the efficiency of gas turbine engines through higher operating temperatures, investment cast equiaxed turbine component parts are now often replaced by components that have either directionally solidified columnar or single crystal (SX) microstructural morphology. During the operation of these SX turbine blades, damage and wear at their leading tip occurs, limiting the lifetime of the SX components and necessitating the replacement of the numerous airfoils within the engines.

Hence, it is of immense interest to develop a process which can restore the single crystal microstructure at the locations of damage and allow for the blades to be reused rather than be scrapped. The work presented here focuses on the development and modeling of a new process, Scanning Laser Epitaxy (SLE), by which high value SX components can be repaired. The SLE process is extensible to metal powder bed- based layer-by-layer additive manufacturing of superalloy components.SLE is a laser-based additive manufacturing process for the creation of structures in equiaxed, directionally solidified and single-crystal nickel superalloys through the selective melting and re- solidification of superalloy powders. In SLE, a tightly focused laser beam is guided by a high- speed galvanometer scanners allowing for tight control of the amount of energy being applied to the top of the powder bed, as well as the speed at which the melt pool moves across the substrate.

Under the proper operating conditions and with sufficient substrate meltback, the solidification microstructure follows the morphology of the underlying substrate, allowing for directional and even SX growth.785Several reports on the modeling aspects of laser cladding-based additive manufacturing approaches can be found in the literature. Approaches based on Finite Element Methods (FEM) and Finite Volume Methods (FVM) are employed for the thermal modeling of the system [1, 2]. Gaumann coupled the Columnar to Equiaxed Transition (CET) modeling with the Rosenthal solution in earlier work based on laser cladding [3]. Rappaz et. al. provided a detailed modeling of the microstructure/columnar orientation based on the solidification velocity and crystalline orientation [4]. This approach is based on the selection of the growth direction that closely follows the melt pool normal direction.

This approach is based on the selection of the growth direction that closely follows the melt pool normal direction. Later on, substrate orientation is also taken into account [5, 6].In the current work, the laser beam moving across the x and y-directions is modeled as a line source scanning in x-direction. The power profile in the y-direction is formulated from the transient averaged data for a single scan and fitted with a 10th degree polynomial. The complete scan pattern is formulated. The associated melting and re-solidification of CMSX-4 powder bed and substrate is formulated by the equilibrium phase change model. Heat transfer to the base plate is also modeled using the heat conduction equation. Analysis of the flow-field modeling reveals the formation of rotational vortices due to the moving heat source and Boussinesq convection. The final melt pool information is used to predict the Oriented to Misoriented Transition (OMT) and CET.

The final melt pool information is used to predict the Oriented to Misoriented Transition (OMT) and CET. The complete simulation along with post-processing to predict the microstructure is performed in the ANSYS CFX modeling environment, which allows the integration of the thermal modeling with microstructure prediction. Reasonable agreement between the simulation and the experimental results is achieved.ExperimentalThe nickel-based superalloy CMSX-4 was chosen as the material to be used in this study of the SLE process. The CMSX-4 powder (composition listed in Table 1), was produced by Praxair Surface Technologies using an atomization process and had a particle size ranging from 85-150µm. Initially, the process was tested on rectangular SX cast CMSX-4 coupons having dimensions of 35.56mm x 10.16mm x 2.54mm. Each substrate was placed into a 35.56mm x 10.16mm recess cut into an Inconel 625 base plate.Table I.

Each substrate was placed into a 35.56mm x 10.16mm recess cut into an Inconel 625 base plate.Table I. Chemical Compositions (wt%) of CMSX-4Cr 7.6Co 9.3Mo 0.4Re 1.0W 2.0Al 12.6Ti 1.3Ta 2.2Hf 0.03Ni BalDuring preliminary testing the powder was deposited in various thicknesses, between 1 and 2mm, to find the ideal height of deposit. The powder was held in place above the substrate using wells cut into a ceramic foam. After the samples were prepared, they were placed into an atmospheric process chamber which was then purged with argon. A 1064nm 2kW Nd:YAG laser beam was used in conjunction with a Cambridge Technologies galvanometer scanner to focus the beam to 1.5mm diameter at the top of the substrate. A raster scan pattern across the width of the sample was used to propagate a linear melt pool across each substrate.

A raster scan pattern across the width of the sample was used to propagate a linear melt pool across each substrate. Three variables were altered between each experimental run: laser power, scan speed, and the number of repeated786scans at the start of the raster pattern. Table II lists the operating ranges for each of the parameters.Table II. Process parameter ranges found to produce well formed CMSX-4 deposits. Process Parameter Powder Thickness Laser Power Scan Speed Initial Repeat ScansRange 1 – 1.5 mm 400 – 600 W 150 – 250 mm/sec 50 – 150The laser power and scan speed were both varied to control the energy being applied to the sample during the raster scan pattern. Sufficient preheat was required in order to form a proper fusion bond across the whole substrate.

Sufficient preheat was required in order to form a proper fusion bond across the whole substrate. Without a high enough preheat the melt pool would not wet the surface of the underlying substrate and voids were formed along the interface.Microstructural InvestigationEach sample was sectioned along the length and width to visualize the microstructure in orthogonal directions. Using a Buehler automated saw, each sample was first cut lengthwise, and then sectioned by multiple widthwise cuts. Imaging was then completed using a Leica DM6000 optical microscope at 100x magnification.As demonstrated by the sample in Figure 1, many sections showed a full metallographic bond along the entire length of the sample. The transition from the cast CMSX-4 substrate to the laser processed material is indicated by the sharp reduction in dendrite spacing.Figure 1.

The transition from the cast CMSX-4 substrate to the laser processed material is indicated by the sharp reduction in dendrite spacing.Figure 1. Representative lengthwise section of the first half of a CMSX-4 sample with the starting edge on the left side of the imageFigure 2(a) provides a detailed view of the transition from the columnar morphology to equiaxed/polycrystalline morphology, known as the columnar to equiaxed transition (CET). This columnar to equiaxed transition (CET) is due to the increase in isotherm velocity near the surface and a decrease in the temperature gradient which causes the formation of a constitutionally undercooled zone [3]. Figure 2(b) provides a detailed view of the transition from single crystal growth with [001] directionality to a shift in orientation.

Figure 2(b) provides a detailed view of the transition from single crystal growth with [001] directionality to a shift in orientation. This change is known as oriented to misoriented transition (OMT) and is caused by a change in direction of the temperature gradient as the substrate becomes increasingly hot toward the end of the scan.787(a)(b)Figure 2. (a) Columnar to equiaxed transition and (b) Oriented to Misoriented transitionThermal ModelingA transient, multi-domain model with temperature dependent property values is used to simulate the thermal aspect of the model and to predict the melt depth. Thermo-physical property data for liquid and solid CMSX-4 is accurately modeled using tabular data [7]. The multi-domain model consists of CMSX-4 domain (Consisting of the substrate as well as the powder deposit), the excess powder surrounding the melt pool, and the Inconel 625 base plate. The Inconel 625 property data is also modeled using tabular data [8].

The Inconel 625 property data is also modeled using tabular data [8]. Figure 3 shows the different domains involved. The mesh for the given problem is generated in ANSYS Mesher. The node count for the given case is around 250k with skewness not exceeding 0.5. The problem is solved in finite volume-based CFD solver ANSYS CFX.Excess PowderCMSX-4 DomainInconel DomainFigure 3. CFX multi-domain model showing different domains in CFD model788The solid and liquid component enthalpy equations, expressed in terms of volume fraction, r, are:( hr SS t S)(hUr SSSS)(kr SST S)I(1a)( hr  LLL t )( hUr LL LL)(kr LLTL)I(1b)Here ρ is the density, h is the enthalpy (including the latent heat), T is temperature, U is velocity and I is the interphase heat transfer. The subscripts S and L refer to the solid and liquid components respectively.

The subscripts S and L refer to the solid and liquid components respectively. Equation (1a) and (1b) can also be expressed more compactly as:( h  t )( hU)Tk ()( hUUY ()(SSLh S))+ S (2)Where (cid:2025)̅ is mixture density, h is the mixture enthalpy, (cid:1863)(cid:3364)		is mixture conductivity, YS is the mass fraction of solid [9].The final term of Equation (2) represents an additional source due to the difference in velocity between the solid and liquid components. This term is finite only in the mushy region of the flow (0< YS, YL <1). In regions of pure liquid or pure solid, the mixture enthalpy equation is identical to the appropriate component enthalpy equation. The volume fraction of liquid and solid component is incorporated using accurate tabular data representing solid fraction of CMSX-4 in the melting range. In order to consider the resistance to flow in the mushy region a Darcy like source term is incorporated.

(cid:4666)K(cid:1487)T(cid:4667) (cid:3397) S(cid:2868) (5)Where, (cid:2025) =Density, H=Enthalpy, K=Thermal conductivity, S0=Source/sink. Since there is no flow into or out of the domain, the chosen model prevents the use of an explicitly varying density for liquid CMSX-4. Therefore the volumetric changes are accounted for using the Boussinesq approximation to artificially include a buoyancy induced source term.789The actual laser heat source has a Gaussian beam radius of 0.75 mm. The spatial profile in the y- direction as the laser beam completes one scan is averaged over a single scan length over the sample width and fitted with a 10th degree polynomial. This particular profile is then used as a line source with Gaussian distribution applied in x-direction. This constitutes the term S in the equation (2). The conservation of H-energy is assured at all connecting interfaces.

This constitutes the term S in the equation (2). The conservation of H-energy is assured at all connecting interfaces. The convection heat loss is modeled using Newton’s law of cooling as a surface sink term applied at all interfaces open to surroundings. Radiation loss is also modeled from open surfaces using sink term. The transient model is solved using a timestep value equal to 0.00215 s. The inclusion of convection, phase change model and Darcy sink tem together necessitated the use of a lower timestep value in order to ensure the solver stability. The imbalance in the system is kept within 1% and the convergence ensures the RMS value of residual to below 1e-4.Once solved, the result is post-processed in CFD-post to analyze the melt pool nature, the melt depth profile and the temperature distribution. The convection modeling allowed visualizing the velocity and vorticity field.

The convection modeling allowed visualizing the velocity and vorticity field. For comparison purposes, pure conduction modes are also run with same set of parameters and the results are compared to distinguish the effect of convection on melt depth and microstructure.Results and DiscussionThe melt depth profile is of particular importance for the present application and hence compared with the actual micrograph. Reasonable quantitative agreement is achieved in the simulated profile and observed data as shown in Figure 4(a) and Figure 4(b) for two different powder thickness and operating parameter sets.(a)(b)Figure 4. Comparison of the simulated meltdepth (Red line) with actual micrograph for (a) powder thickness = 1 mm and (b) powder thickness = 1.5 mm In order to understand the effect of convection on the estimated melt depth, simulations without incorporating fluid convection are carried out and comparisons are drawn with actual melt depth790profile.

As seen in Figure 5, the discrepancy is centered at the starting edge and begins to dissipate as the scan progresses. However, it is found that at any position of the heat source incorporation of convection results in a decrease of melt pool of area by around 10%.Figure 5. Comparison of the melt depth with actual micrograph for models without convection (red line) and with convection (blue line)The effect of convection on the mushy zone is also studied and compared in Figure 6. As shown, the extent of the mushy zone decreases by around 30% due to the inclusion of convection.(a)(b)Figure 6. Extent of the mushy zone (a) without convection and (b) with convectionHowever, the most significant effect of convection is observed on the temperature gradient profile at the solid-liquid interface.

As shown in Figures 6, the melt pool takes a convex hull shape in the raster scan scenario instead of an ellipsoidal profile, as has been observed by others in single bead scans of laser cladding. As seen from Figure 7, inclusion of convection results in a decrease in the range of temperature gradients over the melt pool. As a result, the maximum temperature gradient decreases and the minimum value observed increases. However, the maximum value is always observed ahead of the heat source, in a region which will undergo remelting as the scan progresses. The region that lies behind the heat source is of importance and this region experiences increased temperature gradient due to inclusion of convection.791(a)(b)Figure 7. Temperature gradient at the solid liquid interface (a) without convection and (b) with convectionThe velocity field generated due to the moving heat source is of importance to understand how loose powders are drawn into the melt pool.

As shown in Figure 8 (a), the flow field generated consists of to two vortices with the center of the vortex at lower velocity. The center is located at the middle of the melt pool and the flow field draws loose powder at the top while the melt pool expands at the bottom and sideways. The moving heat source also results in an outgoing high temperature current at lower velocity as shown in Figure 8 (b).(a) )(b)Figure 8. (a) Velocity vector observed in the liquid region and (b) Temperature map of the velocity vector in the melt pool.The presence of the curl component of the velocity vector in Figure 9 indicates that the flow field comprises of rotational vortices in the given case.Figure 9. Vorticity vector in the melt pool792Microstructure ModelingModeling of solidification microstructure is conducted particularly to understand the following two aspects of SX alloys – (a) OMT and (b) CET.(a) OMT Modeling:1.

A value of zero indicates the possibility of conflicting columnar orientations.The plot of the orientation vector is shown in Figure 10 for two different operating parameter sets. Clearly, the flip in direction (from 001 to 100) is suggested in the plot as evidenced in the lengthwise section of the samples. However, Due to the approximation of power profile in the y- direction, the 010 orientation cannot be described.(a)(b)Figure 10. Orientation vector plot showing OMT for two different operating parameters.(b) CET Modeling:The CET modeling is done in the following steps:1. The Temperature gradient at the solid-liquid interface (G) is evaluated from CFD model. 2. The solid-liquid interface orientation (θi) is evaluated from θi =cos-1(Gi/G) (for i=x, y, z) 3. The growth rate of the solid-liquid interface is given by V= S cos θ (S= scan speed/Line source speed)7934.

The growth rate of the solid-liquid interface is given by V= S cos θ (S= scan speed/Line source speed)7934. the temperature gradient parallel to the dendrite growth direction, Ghkl, is calculated using the equation Ghkl = G/cos ψ;5. The dendrite growth velocity, Vhkl, is calculated using the equation Vhkl = Scosθ/ cos (cid:2032)(cid:3038). 6. The Rappaz modification [4] is applied to predict CET as follows:nGhkl/ Vhkl 34 3ln(1)N0 n 1(1 n1Tn Ttipn1 )TtipWhere, n=material constant (n=3.4 for CMSX-4), φ=equiaxed fraction (critical value = 0.066%), N0= nucleation density, ∆Ttip=tip undercooling, ∆Tn= nucleation undercooling.7. A lower value of the left hand side increases φ for a given nucleation density. Hence, a lower value of this ratio indicates the likeliest location of columnar-to-equiaxed transition.n hkl VG /ratio for the same sample at two different positions of Figure 11 shows the plot of the heat source.

Here only the portion of the melt pool lying behind the heat source is compared. As seen, the value varies by five orders of magnitude with position. This explains the random nature of the stray grain formation for this sample as seen. The CET parameter is lower for the 2nd position and the experimental micrograph in Figure 12 also shows CET at lower height for position 2.hkl(a)(b)Figure 11. CET criterion plot for same sample at (a) heat source position 1=7.9 mm and (b) heat source position 2=10.1 mm(a)(b)Figure 12. (a) Experimental micrograph for the given sample showing CET and (b) closer look at the CET for the given position showing larger columnar height for position 1.794ConclusionIn the current work, a modeling scheme has been implemented to simulate the SLE process using a single platform. The model is able to predict the complete temperature distribution and melt pool behavior.

The model is able to predict the complete temperature distribution and melt pool behavior. Inclusion of convection modeling provides insights on the flow field established in the melt pool and how the loose powders are drawn inside melt pool. It is shown that the flow field comprises of two rotational vortices. The same platform also allows the modeling of solidification microstructure to predict the OMT and CET. The melt pool behavior for raster scanning situation is investigated. The final aim would be to develop a processing map based on the CET criteria that will allow us to predict the microstructure beforehand. However, the current simulation still allows us to find out the possibility of getting CET and OMT as well as the most likely locations of their occurrence. The simulation shows reasonable qualitative agreement with the experimental data in predicting the CET.

The simulation shows reasonable qualitative agreement with the experimental data in predicting the CET. However, the inclusion of Marangoni convection and the determination of CET criterion from experimental studies is necessary to obtain quantitative data for microstructural investigation e.g. columnar to equiaxed ratio.AcknowledgmentsThis work is sponsored by the Office of Naval Research through grant N00014-11-1-0670.References[1][2]T. D. Anderson, J. N. DuPont, and T. DebRoy, "Origin of stray grain formation in single- crystal superalloy weld pools from heat transfer and fluid flow modeling," Acta Materialia, vol. 58, pp. 1441-1454, 2010. T. H. C. Childs, C. Hauser, and M. Badrossamay, "Mapping and Modelling Single Scan Track Formation in Direct Metal Selective Laser Melting," CIRP Annals - Manufacturing Technology, vol. 53, pp. 191-194, 2004.[3] M. Gaumann, "Epitaxial Laser Metal Forming of a Single Crystal Superalloy," PhDThesis, EPFL Lausanne, 1999.[4] M.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCOMPUTATIONAL MODELING OF THE INERT GAS FLOW BEHAVIOR ON SPATTER DISTRIBUTION IN SELECTIVE LASER MELTINGXiaobing Zhang1,2, Charles Tuffile2, Bo Cheng2,*1Department of Mechanical and Aerospace Engineering Rutgers, The State University of New Jersey Piscataway, NJ, 08854, USA 2Robert Bosch LLC, Cambridge, MA, 02139, USA *Corresponding author: Bo.Cheng@us.bosch.comAbstractIn the selective laser melting (SLM) process, the spatter, introduced by moving melt pools, can contaminate the build region and potentially reduce the part quality. A well-designed inert gas flow system can help to effectively remove the process emissions. Therefore, it is critical to understand the flow characteristics as well as the interactions between gas flow and spatter.

Therefore, it is critical to understand the flow characteristics as well as the interactions between gas flow and spatter. In this study, a Computational Fluid Dynamics (CFD) model for the SLM gas flow system has been developed to simulate the complicated flow behavior in the chamber. The generation of spatter over the build region has been simulated by a discrete phase model (DPM). The Eulerian- Lagrangian modeling technique has been applied to accurately capture the influence of gas flow on ejected spatter. The process variables that can significant affect the gas flow efficiency in the chamber, such as inlet design, volume flow rate and material property, have been numerically investigated.Keywords: Selective Laser Melting (SLM); Computational Fluid Dynamics (CFD); Discrete phase method (DPM); Gas flow; Spatter1. IntroductionIn selective laser melting (SLM) process, a high-energy laser beam is utilized to melt and fuse the metallic powder materials.

IntroductionIn selective laser melting (SLM) process, a high-energy laser beam is utilized to melt and fuse the metallic powder materials. The accompanied high local temperature can easily exceed the material evaporation point and cause evaporation. The complicated vaporization effect may cause a “vapor-jet” effect, leading to the generation of emissions. The emissions include the vaporization gases, “vapor jet” entrained powders, and liquid droplets ejected from melt pool due to melt pool dynamics. Therefore, the spatter (ejected particles from or around melt pool) mainly consists of droplet spatter and powder spatter [1]. Numerical Modeling has been utilized to understand the complex process physics in powder bed additive manufacturing process. Lee and Zhang [2] developed a discrete element method (DEM) model to simulate the powder spreading process in SLM.

Lee and Zhang [2] developed a discrete element method (DEM) model to simulate the powder spreading process in SLM. In addition, a computational fluid dynamics (CFD) model is developed to investigate the melt pool dynamics with the incorporation of DEM generated powder bed. The thermal information was utilized to calculate solidification morphology and grain size. Panwisawas et al [3] simulated the positions of powder particles in a powder bed by dropping spheres based on encountered obstacles. Micro-scale CFD simulations were performed to establish the relationship between laser process parameters, melt pool and microstructure. Shrestha and Chou [4] modeled the two-layer powder deposition process in SLM, their simulation showed interlayer lack-of-fusion1362phenomenon if a high laser scanning speed was applied. Using both DEM and CFD modeling tool, Cheng et al [5] studied the powder spreading condition on melt pool characteristics in SLM process.

Using both DEM and CFD modeling tool, Cheng et al [5] studied the powder spreading condition on melt pool characteristics in SLM process. A higher powder bed packing density, caused by smaller roller rotational speed, lead to smaller re-melting depth. Cheng et al [6] also investigated the evolution of porosity in a single track scanning process. Pores could be formed by the collapse of keyhole. Lawrence Livermore National Laboratory (LLNL) utilized the arbitrary lagrangian-eulerian (ALE) package to study various physical processes such as melt pool dynamics, scanning track surface morphology, and melt pool emission [7, 8, 9, 10]. A comprehensive review of additive manufacturing process thermal analysis and simulation can be found in literature [11]. However, it is very difficult to simulate the spatter motion over the entire build chamber using the micro-scale melt pool model.On the other hand, spatter characteristics can be studied by experimental measurement.

Using the high-speed X-ray imaging technology, Zhao et al [12] were able to obtained important spatter parameters such as spatter count, diameter, and velocity with respect to different scanning speeds and laser powers. Their observation indicated that more and larger speed spatters could be generated from higher laser power condition. Gunenthiram et al [13] also experimentally analyzed the formation of spatters in SLM process. By introducing a high speed imager in SLM system, they quantified the spatter ejections as well as possible correlation with melt pool dynamics. Their results showed that larger sized spatters could be generated for SS316L material if high laser volume energy density was applied on powder bed.

Their results showed that larger sized spatters could be generated for SS316L material if high laser volume energy density was applied on powder bed. On the other hand, the droplet might be incorporated into melt pool area for Al-12Si material, thus higher spattering was observed on SS316L powder bed.Without any removal process, the spatters may re-deposit on the powder bed due to gravity and contaminate the build area, potentially affecting the build part quality. In common practice, a shielding gas flow can help to remove the ejected particles, as shown in Figure 1. However, the spatter removal efficiency heavily depends on the design and optimization of the gas flow system in the build chamber. Studies about the gas flow system are sparsely found in the literature. Philo et al [14,15] developed a CFD model to study the gas flow characteristics inside the Renishaw AM250 build chamber. The movement of spatters was simulated by discrete phase method (DPM).

Philo et al [14,15] developed a CFD model to study the gas flow characteristics inside the Renishaw AM250 build chamber. The movement of spatters was simulated by discrete phase method (DPM). The simulated velocity field in the chamber was validated against the flow field obtained by Hot Wire Anemometer (HWA) measurement. They found that the flow uniformity and average flow velocity could be influenced by inlet rail radius and inlet/outlet height. Chen et al [16] implemented a CFD tool to optimize the gas flow inside Renishaw AM250 chamber. A secondary downward flow from the to chamber, besides the main horizontal inlet flow, was critical to avoid the recirculation of gas flow inside build chamber. Wang and Chang [17] tried to optimize the flow uniformity inside a selective laser sintering (SLS) system.

Wang and Chang [17] tried to optimize the flow uniformity inside a selective laser sintering (SLS) system. A CFD model was utilized to investigate the effect of influential factors on flow velocity profile and uniformity, e.g., the trapezoid push nozzle, the suction tunnel and nozzle-to-plane distance. However, optimal design and modeling for build chamber to increase spatter removal rate, especially considering gas flow – spatter interaction, is still not thoroughly investigated.1363Build ChamberOUTLE ~Inert gas•INLETPowder bedFigure 1. Graphic description of problem.This study developed a comprehensive fluid-particle interaction CFD model to simulate the gas flow effect on spatter distribution. The spatter motion trajectory is studied to better understand the particle dynamics inside the gas flow chamber.

The spatter motion trajectory is studied to better understand the particle dynamics inside the gas flow chamber. Several influential process and design parameters of the gas flow system, such as flow rate, inlet height and width, as well as material density, have been investigated to enhance the removability of the process emissions.2. Numerical MethodologyIn this study, the objective is to maximize the removal efficiency of process emissions during the SLM process. Therefore, the design of the gas flow system, including gas inlet width, inlet height, volume flow rate, and processing materials, will be thoroughly investigated. A generic baseline SLM build chamber design (Renishaw AM250) is shown in Figure 2 (a). The commercial simulation package, ANSYS Fluent, was used to build a gas-spatter interaction model. A CFD fluid flow model was used to describe the gas flow characteristics inside build chamber. The fluid flow is assumed to be steady, incompressible and turbulent flow.

A CFD fluid flow model was used to describe the gas flow characteristics inside build chamber. The fluid flow is assumed to be steady, incompressible and turbulent flow. Standard Navier-Stokes equations and k-e turbulent model are used to describe the fluid characteristics [14,15].In addition, a Discrete Phase Method (DPM) model was developed to study the spatter distribution within the SLM chamber. Spatters are generated by DPM method and fully coupled with the gas phase.

Spatters are generated by DPM method and fully coupled with the gas phase. To efficiently model the gas-spatter interaction inside the chamber, several assumptions have been made:(1) Instead of generating the spatters from the melt pool level, the spatter source are modeled as five lines evenly distributed on the powder bed to cover the entire build area, as shown in Figure 2 (a).(2) The spatter particles are ejected with three different ejection angles to the horizontalsubstrate: 60(cid:953), 90(cid:953), and 120(cid:953).(3) All the ejection sources have Gaussian distributed particles ranging from diameter 20(cid:586)(cid:143) to 80 (cid:586)(cid:143) with an ejecting speed of 3 m/s.(4) The spatter material is SS316L with a density of 7950 kg/m3.Different boundary conditions have been applied to different domains of the model. To calculate the spatter-chamber wall surface contact, “Wall” boundary condition with DPM settings are used.

To calculate the spatter-chamber wall surface contact, “Wall” boundary condition with DPM settings are used. The DPM particles will be rebounded off from the “reflect” wall boundaries with a1364modified momentum. The trajectory calculations are terminated and recorded as “trapped” if the particles fall on “trap” wall boundaries. In this simulation, only the powder bed surfaces, including the five spatter source lines, are listed as DPM “trap” boundary conditions. Other chamber wall surfaces are marked as “reflect” boundaries. The inert gas firstly enters the inlet rail with given flow rate, then flows into the main build chamber through a row of cylindrical nozzles with 12 mm in diameter. The gas and spatter flow out of the domain through outlet. Therefore, the gas inlet is set to be “mass flow rate” while “outflow” boundary condition is applied for the outlet. The inert gas used in the simulation is argon with a viscosity of 0.00002125 kg/m-s and a density of 1.6228 kg/m3.

The inert gas used in the simulation is argon with a viscosity of 0.00002125 kg/m-s and a density of 1.6228 kg/m3. The mass flow rate is set as 0.00677 kg/s (250 L/min).The entire domain was meshed using Tetrahedron elements. To capture flow characteristics at boundary layer, 8 inflation layers were added upon the substrate. Figure 2 (b) shows the model mesh information, the model has ~1.5 million elements. For all simulated cases in this study, there may be small variations in the total amount of element which is caused by the slight change of gas chamber geometry.Inlet designSpatter source'---~ = =-----:~ =='>o.•1 ... ,---===0•,1---====:::::,0,2 (m)0.150.05(a) Geometry of gas chamber(b) Mesh used for computational domain Figure 2. Model geometry and mesh information.To solve the coupled fluid-particle momentum and pressure equations, the SIMPLE (Semi- Implicit Method for Pressure Linked Equations) solution method was used.

Model geometry and mesh information.To solve the coupled fluid-particle momentum and pressure equations, the SIMPLE (Semi- Implicit Method for Pressure Linked Equations) solution method was used. The solution was considered to be converged when the residuals of continuity and momentum were smaller than 1×10−4. High-performance computing (HPC) system is utilized to conduct simulations. The1365specification of the computing node is: 2 × Intel Haswell E5-2680v3 (24 cores per node) and 64 GB RAM. It takes ~2 hours to obtain converged results using two nodes.3. Results and DiscussionIn this part, one baseline design simulation is used to show the gas flow variations and particle behaviors. Afterwards, three different design parameters, inlet width, inlet height and flow rate, are adjusted to show their influence on the system and spatter removal. Finally, different spatter material, namely, AlSi10Mg, Ti64 and SS316L, have been simulated to study the particle density effect.

Finally, different spatter material, namely, AlSi10Mg, Ti64 and SS316L, have been simulated to study the particle density effect. To quantify emission removal performance, a “clear rate” is used to describe the spatter removal efficiency. The clear rate is defined as the ratio of the number of spatter particles removed out of the build region (powder bed) to the total number of spatters.3.1 Overview of the baseline modelFigure 3 shows the simulation results for gas-spatter interaction in the baseline design chamber. The spatters firstly ejected to the gas chamber from the build region, then the incoming gas flow blew the particles toward the outlet. However, not all spatters can be removed from the build chamber, they may re-deposited to the chamber substrate due to gravity.Velocity1 .500e+0001 .125e+0007 .S00e-0013 .750e-0010 .000e+000[m s"-1](a) Blow away(b) Re-depositionFigure 3.

Typical simulation results.Two perspectives of the velocity contours inside the system are discussed below. Figure 4 (a) is a side view of the chamber gas flow velocity contours located in the middle of the chamber. It is seen that high gas velocity streams come out of the inlet nozzles and flow down toward the substrate. The spatter trajectories are affected by the gas flow and spatters move toward the outlet, as shown in Figure 4 (b). It is seen that the spatter generated in the region closest to the outlet are mostly removed. Figure 4 (c) is a top view of the gas flow velocity contours on the center-plane (center of inlet nozzle), where a high velocity concentration region is observed around the inlet region with a near triangular shape. The spatter distribution map shows that the substrate region close to inlet has little spatter concentration while the region close to outlet has high spatter concentration, as shown in Figure 4 (d).

The spatter distribution map shows that the substrate region close to inlet has little spatter concentration while the region close to outlet has high spatter concentration, as shown in Figure 4 (d). The spatter re-deposition map generally follow the velocity contour: high velocity region has little re-deposition phenomenon. With the decrease of the inlet gas flow velocity, the spatters fall down to substrate region that is away from the inlet region.

Velocity contours and spatter distribution of the baseline chamber design.3.2 Inlet flow rate effectThe movement of the spatter is largely decided by gas flow and particle condition. Therefore, a strong inlet gas flow may help to improve spatter clear rate. Three different inlet flow rates, 0 L/min, 250 L/min and 500 L/min, were tested to investigate flow rate effect on spatter distribution. A flow rate of 0 L/min means no gas flow is applied. The spatters simply eject outward then fall back under the gravity force. The drag force imposed on the particles is little due to the negligible gas flow rate. With the increase of flow rates, the gas flow starts to affect the particle motion and the spatter trajectory changes significantly, e.g., the spatters hardly reach outlet for 0 L/min case while majority of the spatters can flow through the outlet for 500 L/min case, as shown in Figure 5.

The clear rate for the three cases are: 0%, 60.4%, 81.1% for 0 L/min, 250 L/min, and 500 L/min respectively. Therefore, high flow rate is helpful for spatter removal.1.8Q0e.0151 600e-0151.400e-0151.2009-0151.000.0158.000e-0166.000.0164 000.0162.000.016~,~10.000e+OOO1367(a) 0 L/min (b) 250 L/min (c) 500 L/min Figure 5. Spatter re-deposition maps and trajectories with different flow rates.3.3 Inlet width effectThe inlet width has been varied proportionally to study the inlet width effect, all other parameters are kept unchanged as the baseline case, e.g., a flow rate of 250 L/min is used. The nozzle to nozzle distance has been adjusted accordingly. The effects of three inlet widths of 225 mm, 250 mm and 275 mm have been compared, as is shown in Figure 6. The flow field differences between different inlet width cases are subtle.

The effects of three inlet widths of 225 mm, 250 mm and 275 mm have been compared, as is shown in Figure 6. The flow field differences between different inlet width cases are subtle. In all cases, the flow velocity is stronger in the middle of the chamber, but this effect is particularly noticeable for a smaller inlet width. Larger inlet widths such as 275 mm, results in a lower gas flow velocity above the build area which diminishes the particle removal efficiency.

Larger inlet widths such as 275 mm, results in a lower gas flow velocity above the build area which diminishes the particle removal efficiency. An inlet width with same or similar width as the build region (powder bed) width of 250 mm is shown to be most effective, as the highest particle clear rate of 60.4% with the inlet width of 250 mm, compared to the rate of 60.0% for a narrower width of 225 mm and 59.3% for a wider width of 275 mm.Particle mass concentration2.~ 1 51.350e+OOO1.200e+OOO1.0SOe+OOO9 .000e-0017 .SOOe-OO 16.000e-0014.SOOe-001 3 .OOOe-OO 1 1.SOOe-OO 1O.OOOe+OOO(ms"-1)(a) Width = 225 mm (b) Width = 250 mm(c) Width = 275 mmFigure 6. Spatter re-deposition maps and velocity contours with different inlet widths.3.4 Inlet height effectThe inlet height effect on spatter re-distribution is also investigated. Inlet height is defined as the distance from the center plane of inlet nozzles to the substrate plane.

Inlet height is defined as the distance from the center plane of inlet nozzles to the substrate plane. Three designs with different inlet heights: 50 mm, 67.5 mm and 85 mm are compared in this section while all other parameters are the same as the baseline case. Figure 7 shows the re-deposition maps and velocity contours at center vertical plane. It is found that the case with inlet height equal to 67.5 mm (Figure 7 (b)) removes the most particle emissions (60.4% clear rate). Lower inlet height design (Figure 7 (a)) has a clear rate of 59.8%, and higher inlet height design (Figure 7 (c)) has a clear rate of 57.7%. It demonstrates that a moderate inlet height is preferred to remove the most particles. It is observed that the high gas velocity region moves up and down along with the inlet height modifications. The spatters have pre-defined initial ejection velocities, and it falls after reaching a peak height due to gravity.

The spatters have pre-defined initial ejection velocities, and it falls after reaching a peak height due to gravity. Therefore, spatters may penetrate the high velocity region with a low inlet height1368design. On the other hand, the high gas velocity region may be too high to exhibit influence on the spatter particles for a high inlet height design case.Particle mass concentration2J)OOe--015UOOe--015, __0,000e+OOO(kg m"-31 Velocity1.500&+0001.350e+OOO1.0508+0009.000e-OO,7.SO()e.()()16.000e-OO,4.500e--001 3.oooa-oo, 1.500e--0010.000.,000[ms"-1I(a) Height = 50 mm (b) Height = 67.5 mm (c) Height = 85 mmFigure 7. Spatter re-deposition maps and velocity contours with different inlet heights.3.5 Spatter material property effectTo examine the effects of the spatter material property, three materials are chosen and compared using the baseline case. The three common SLM materials are AlSi10Mg, Ti64, and SS316L.

The three common SLM materials are AlSi10Mg, Ti64, and SS316L. The corresponding densities are about 2700 kg/m3, 4400 kg/m3 and 7950 kg/m3. Even though the interactions between the particle and gas fluid are considered in the simulations, the influence of the particle on the gas fluid field is minor such that the velocity fields remain almost the same for these three cases. Figure 8 shows the spatter particle trajectories of the three materials. It is shown that the spatter material is also a significant factor for the system design. The particle trajectory with lighter material shows higher similarity to the gas flow velocity field. This can be explained that the ratio between the flow drag force and the gravity will be larger for a lower density spatter material, therefore, the individual particle motion is dominated by fluid drag force.

The resulted clear rate is higher for lighter materials, e.g., 69.4% for AlSi10Mg, 68.3% for Ti64 and 60.4% for SS316L.Particle mass concentration2.000e--0151.800e-0151.600e--0151.400e-015 1.200e.Q151.000e~)158.000e--0166.000e--0164.000&-0162.000e--016 O.OOOe>OOO!kgm"-311369(a) AlSi10Mg (b) Ti64 (c) SS316. Figure 8. Spatter re-deposition maps and trajectories with different processing materials.4. ConclusionIn this study, a fully coupled CFD-DPM model has been developed to investigate the gas flow effect on spatter distribution inside the SLM build chamber. The fluid flow characteristics have been calculated by a steady-state and turbulent CFD model. In addition, the spatter generation is modeled by DPM model. The influence of different build chamber parameters, such as inlet flow rate, inlet width and height, and spatter materials, have been studied.

The influence of different build chamber parameters, such as inlet flow rate, inlet width and height, and spatter materials, have been studied. The major findings can be summarized as follows:(1) The flow rate has considerable effect on the spatter particle removing process. It is observed that a higher flow rate can significantly affect the spatter moving trajectory toward the outlet, which can contribute to obtain a higher clear rate, e.g., a clear rate of 81.1% for a flow rate 500 L/min while it is 60.4% for a flow rate of 250 L/min.(2) Three inlet widths were compared using the developed model. It is found that the inlet width has minor effects on the clear rate. An inlet width that is similar to the width of the build area (powder bed region) shows best performance for clear rate.(3) The effect of inlet height has been investigated as well.

An inlet width that is similar to the width of the build area (powder bed region) shows best performance for clear rate.(3) The effect of inlet height has been investigated as well. Spatters may penetrate the high velocity region (low inlet height design), or the high gas velocity region may be too high to blow away spatters (high inlet height design). A moderate height is preferable for spatter removals.(4) The spatter materials have a noticeable effect on particle motion trajectory and clear rate. A higher similarity to the gas flow velocity stream field is shown for the trajectory of lighter spatters. It is because the spatter motion is dominated by fluid drag force for low density spatter. Therefore, higher clear rate is expected for lighter spatters, e.g., clear rate is 69.4% for AlSi10Mg while it is 60.4% for SS316L.The future work of this study will be focused on comprehensive validation of the developed model.

Models are enhanced by measured thermophysical material properties. Validation of the solidified melt geometry showing that conductive mode melting and instabilities such as balling can be captured with existing models and pave the way for models which capture the onset of keyholing. Examination of the melt track microstructures can also be used to determine local cooling rates, granting insight into the phase evolution differences between the alloys.IntroductionAdditive manufacturing (AM) is deemed a disruptive technology; a technological enabler for industry 4.0 [1], [2]. In particular, laser powder bed fusion (LPBF) has received great interest due to the production of highly functional and complex parts with 99.9%+ densities without the need of additional post processing [3]–[9]. Furthermore, the resulting mechanical properties of LPBF are superior to parts produced by conventional manufacturing techniques [8]–[11].

Furthermore, the resulting mechanical properties of LPBF are superior to parts produced by conventional manufacturing techniques [8]–[11]. This is a result of the unique thermal cycle involving steep thermal gradients and rapid cooling caused by rapid melting and high cooling rates (approx.105-106 oC.s-1), with the potential of re-melting (location dependent) [12]. It is these primary variables that determine the characteristics of an alloy’s final microstructure and the inherent defects that follow. The phenomena experienced presents an opportunity to create alloys with unique material properties that could not be reproduced by more conventional processing techniques such as casting. Such alloys that show great promise include nickel-chromium-based super alloy, Inconel 625 (IN625).IN625 is extensively used across a wide range of industries due to its high-performance capabilities during the application of elevated temperatures and pressures.

Its strength is derived from the stiffening effect of alloy additions, Mo and Nb, within the Ni-Cr matrix. The combination of elements results in superior corrosion resistance properties and high- temperature effects such as carburization and oxidation [13].BackgroundTraditionally, trial-and-error approaches are employed to optimise the LPBF process which results in cost, material and time wastages thereby limiting the industrial application of this technique. To combat this multi-physics modelling techniques are utilised in aid of process and part optimisation [14]–[22]. Although, due to the enormous time- and length-scales1200involved, models are split into micro-, meso- and macroscale processes which include modelling of thermal history, residual stress and material melting/solidification [23].

ESI Group have produced a complete suite of tools capable of simulating the various physical phenomena that occur during the LPBF process to identify manufacturing defects and residual stresses input by this process.the spreading process,The focus of these works attempt to validate the melting and solidification module which uses information obtained via the powder coating module. The powder coating module utilizes a discrete element approach to simulate powder spreading onto a substrate. This can then be enhanced to simulate powder coating on previously consolidated layers, but for the purpose of this work only a single layer is modelled. The resulting geometry can then be imported into the high-fidelity melting module which determines laser-powder interaction, heat and fluid flow in the melt pool. For validation purposes, a single line study is presented in this document.

For validation purposes, a single line study is presented in this document. Several authors have conducted similar strategies to investigate single track formation for parameter optimisation and model validation purposes [24]–[28]. Using various process parameter settings, single lines on top of a substrate are processed. The resulting line morphologies were unique to each specified setting applied. Using ESI-AM, similar lines were processed and compared to the physical result for model validation.Materials and MethodsInconel 625 PowderFor the purpose of this study nitrogen gas atomised Inconel powder, IN625 (prepared by Sandvik Osprey, UK) was used; the chemical composition and powder morphology is shown in Table 1 and Figure 1 respectively. Particle size distribution (PSD) was first analyzed by using laser diffraction analysis (LDA) resulting in a D10, D50 and D90 of 24.5 μm, 35.2 μm and 57.2 μm respectively.

Particle size distribution (PSD) was first analyzed by using laser diffraction analysis (LDA) resulting in a D10, D50 and D90 of 24.5 μm, 35.2 μm and 57.2 μm respectively. The majority of the IN625 powder is spherical in shape with a smooth surface; a typical result of the gas atomisation process. At [x500] magnification some satellite formation can be observed, but this effect was minimal and posed no concern regarding powder rheology.Table 1. Limiting and actual chemical composition of IN625 (wt.%) (prepared by Sandvik Osprey, UK) [13].Ni58 min.Cr20.0- 23.0Mo8.0- 10.0Fe5.0 max.Nb3.15- .15Al0.4 max.C0.1 max.Si0.5 max.Co1.0 max.Mn0.5 max.Ti0.4 max.P 0.015 max.S0.015 max.62.021.78.93.83.740.0070.020.02<0.010.01<0.01<0.003<0.0011 ~I\ a..f ~ . ,:. '"", _I., ,,,(), nO_~l){:)n,,l-J;> __ r,,.L'i'?,Q,~( 8iicJJf c~~o {.,(,~11£>~ C\(}oL~J~) .,f) 'ot.\,lfi(..~c.i.. ..

This process is then repeated throughout the build over multiple layers to form the final part. The equivalent scan speed, V, was then calculated by dividing PD by ET. Argon gas was used to flood the build chamber; reducing the oxygen level to <0.1% and forming a completely inert atmosphere. In addition to this, reduced build volume (RBV) auxiliary equipment was used in conjunction, enabling efficient part production and rapid material change-over.PD-------------------------------- ....Laser DirectionFigure 2. Application of a modulated laser: Heat is applied at a singular point for a specified exposure time, ET, and then moves a specified point distance, PD. A laser spot size/beam diameter, σ, of 70μm is kept constant whereas hatch spacing, HS, was varied.Process Parameter OptimisationDuring the LPBF process, the amount and rate of energy transferred to the powder is governed by the process parameters applied.

In addition to the resulting L9 configurations, a 10th configuration was tested as suggested by Renishaw. The optimised configuration was used to manufacture the substrate, namely the crucible (‘The Crucible Method’ as first described in [30]), of which the single lines were processed on.1202Table 2. Parameter sets applied to each density cube sample to obtain the optimum parameter set for this alloy.

(cid:1831). (cid:1830) =(cid:1842) (cid:1848)=(cid:1842). (cid:1831)(cid:1846) (cid:1842)(cid:1830)(2)Thirty single lines (22.5mm in length) were produced with various L.E.Ds applied to promote a range of melt pool geometries for validation against results produced by the melting model later described. Three power settings were applied; 300 W, 350 W and 400 W. For each power setting 10 exposure times were applied ranging between 23.33 μs and 233.33 μs (or 3.00 m/s to 0.30 m/s). The resulting L.E.Ds varied from 100 J/m to 1333.33 J/m. This is seen in Figure 3. A single power set (10 lines of varying exposure) was processed on top of a ‘crucible’ – a substrate manufactured by LPBF using the optimum settings for this powder.

This is seen in Figure 3. A single power set (10 lines of varying exposure) was processed on top of a ‘crucible’ – a substrate manufactured by LPBF using the optimum settings for this powder. The Crucible Method’s, crucible consists of a 3D tapered rectangular structure as detailed in Figure 4.140012001000)m/ J (800D E L..60040020000.30.60.91.21.51.82.12.42.73Scan Speed (m/s)300W350W400WFigure 3. L.E.Ds applied to lines studied. A combination of 3 power setting and 10 exposure times (effective scan speeds) were applied resulting in a total of 30 parameter sets (i.e. 30 single lines).120325 1· __ · 110..I.Jt ~~49~ .I ~1JLI. "Figure 4.

30 single lines).120325 1· __ · 110..I.Jt ~~49~ .I ~1JLI. "Figure 4. Isometric view (left) and detailed engineering drawing (right, dimensions in mm) of substrate produced by LPBF, namely, the Crucible Method.Metallographic Preparation and Optical AnalysisTop-down images were taken using optical microscopy (Zeiss Smartzoom 5) to assess overall line morphology. Following this, each crucible was machined in half to expose the central cross section of each line. They were then mounted by compression mounting and polished according to Buehler’s recommendation of Ni-based super alloys [31]. For microstructural characterisation, the crucibles were then electro-etched using 2% H2SO4 and H2O under a voltage of 5 V for 5 to 15s. Using optical microscopy (Zeiss Observer), the melt pool of each line was characterised in terms of melt pool width (w), depth (d) and height (h) as depicted in Figure 5.

Using optical microscopy (Zeiss Observer), the melt pool of each line was characterised in terms of melt pool width (w), depth (d) and height (h) as depicted in Figure 5. The melt pool width was defined by the distance of the weld along the level of the substrate. The melt pool depth was defined by the vertical distance from substrate to the bottom of the weld. The melt pool height was defined by the distance from the substrate to the top of the weld. Three sets of measurements were taken by re-grounding, polishing and etching to obtain the average measurement of width, depth and height of each line processed.Figure 5. Techniques used to characterise 3 types of melt pool formation; deep-weld (left), medium-weld (right) and a ‘balled’-weld.

Techniques used to characterise 3 types of melt pool formation; deep-weld (left), medium-weld (right) and a ‘balled’-weld. This type of weld formation is relative to the parameter set applied as described later in this study.Spreading ModelThe powder spreading model is used to analyse powder flowability which is dependent on the system geometry, material density and particle size distribution (PSD) applied. The resulting spread is then imported into the melting model to serve as a numerically produced powder bed. In this case, the spreading geometry (depicted in Figure 6) consisted of a layer thickness (LT) of 60 μm, spreader thickness (ST) of 100 μm, spreader length (LSpreader) of 4 mm, table length (LTable) of 4 mm and table width (w) of 1 mm.1204LTSTwLSpreaderLTableFigure 6. Spreading model setup and parameter definitions.PSD was analyzed by using laser diffraction analysis (LDA) resulting in a D10, D50 and D90 of 24.5μm, 35.2μm and 57.2μm respectively.

Spreading model setup and parameter definitions.PSD was analyzed by using laser diffraction analysis (LDA) resulting in a D10, D50 and D90 of 24.5μm, 35.2μm and 57.2μm respectively. The particles generated in the spreading model were based on 14 data points encompassing the range of particles measured w.r.t mass fraction. A solid density of 8400kg/m3 was also applied.Melting ModelUsing the melting model, single lines were processed using a variety of settings which would result in a range of line L.E.Ds comparative to those applied in the physical experimentation. Initially 6 melting model cases (Table 3, models 1-6) were developed with L.E.Ds ranging from 107.14 J/mm to 1400.0 J/mm. IN625 and Argon gas were selected as the powder and gas materials respectively (material data supplied by ESI Group).

IN625 and Argon gas were selected as the powder and gas materials respectively (material data supplied by ESI Group). Due to the lack of time and computational resources available, a low-quality mesh was applied to models 1-6 with only a total 34,560 cells which encompassed the gas, powder and substrate domains. Due to the resulting cell size applied, the mesh generated failed to capture an accurate geometry of the powder bed. Following that, a 7th model was simulated using a much higher quality mesh (1,684,800 cells) for comparison (Table 3, model 7). The meshes generated are shown in Figure 7. An arbitrary location was selected on the resulting powder bed to apply the melting domain with a length, height and depth of 900, 460 and 400 μm respectively. A time-step of 1E-7s was applied to each case.Table 3. Severn models were simulated with different parameter sets applied to apply various L.E.Ds in a single line model.

A time-step of 1E-7s was applied to each case.Table 3. Severn models were simulated with different parameter sets applied to apply various L.E.Ds in a single line model. Models 1 to 6 were deemed of low quality due to the application of a course mesh, but an addition 7th model was created with the application if a high quality mesh. Throughout this study the point distance and time-step remained constant.ModelL.E.D (J/m)P (W)ET (μs)PD (μm)V (m/s)No. Cells1107.14300252.802600.004001050.673700.003501400.504792.86300185700.383456051000.03502000.3561400.04002450.297228.57400401.7516848001205Figure 7. Mesh applied to models 1-6 (left) and 7 (right). Due to the large cell size applied to models 1-6, the resulting mesh generated failed to capture the powder bed geometry.

Mesh applied to models 1-6 (left) and 7 (right). Due to the large cell size applied to models 1-6, the resulting mesh generated failed to capture the powder bed geometry. By refining the mesh, the geometric curvature was captured resulting in a high-quality powder bed.Results and DiscussionOptimised Process ParametersIn order to obtain the optimum process conditions for alloy IN625, a density study was conducted using various parameter sets produced by implementation of an L9 orthogonal array. In addition to the 9 generated sets, a 10th was added as recommended by the machine developers, Renishaw. Results are shown in Figure 8 below. The results of this study revealed a typical correlation between V.E.D applied and the resulting part density. As expected, maximum porosity was observed during an application of the lowest energy density. The density increases to 99%+ with sample A1, with the exception of sample A9 which was deemed an error reading.

The density increases to 99%+ with sample A1, with the exception of sample A9 which was deemed an error reading. Although the parameter set for A1 were efficient in this case, it was decided that Renishaw’s suggested setting should be used. The optimum parameter sets used to manufacture the crucible (the substrate) are shown in Table 4.Table 4. Optimum parameter set found by means of density measurement analysis based on Archemedes principle. The optimum value is highlighted in red.Optimum Settings (Renishaw Setup)PD (μm)70.00HS (μm)70.00ET (μs)40.00P (W)400.00V (m/s)1.75V.E.D (J/mm3)54.42)%( y t i s n e D100.0095.0090.0085.0080.00 75.00•A9A6A8A1R• A5 A2• A7 A4A370.0030.0035.0040.0045.0050.0055.00 V.E.D (J/mm3)60.0065.0070.0075.0080.00Figure 8. Resulting densities of parameter sets tested. Majority of V.E.Ds applied resulted in 99%+ dense parts.

Resulting densities of parameter sets tested. Majority of V.E.Ds applied resulted in 99%+ dense parts. Optimum result, R, highlighted in red.1206Single Line - Line MorphologiesFigure 9 gives an overview of each line processed on top of the crucible with a specific parameter set applied. From the top line to the bottom, L.E.D is decreased. During the application of a high L.E.D, lines appear to be thick and stable (e.g. 400W, 0.30 to 0.90 m/s). As the energy density is decreased instabilities result in track irregularities which causes balling (e.g. 400, 1.50 to 2.10 m/s). A further decrease resulted in little or no adherence due to insufficient power and/or exposure time to melt the powder (e.g. 300W, 2.40 to 3.00 m/s).In LPBF process parameters that result in stable and continuous tracks are desirable. Track irregularities can cause various defects such as porosity and may lead to part failure, it is therefore essential to provide enough energy to adequately melt the powder.

Track irregularities can cause various defects such as porosity and may lead to part failure, it is therefore essential to provide enough energy to adequately melt the powder. Conversely, it is also evident that the spatter produced during processing is a function of energy density applied, this may also affect the building process and final part. This furthers the importance of parameter optimisation.0.300.600.90) s /1.20m(d e e p S n a c S1.501.802.102.402.703.00300350 Power (W)400Figure 9. Top-down micrograph showing single line tracks as a function of scan speed (m/s) and power (W). Both layer thickness and spot size remained constant at 60 and 70 μm respectively.Single Line - Cross Section MorphologyTrack morphologies are shown in Figure 10 and the measured widths, depths and heights are shown in Figure 11, Figure 12 and Figure 13 respectively.

Three repeat cross- sectional measurements were taken for each track and the standard deviation is shown. Typically, application of a high line energy density resulted in stable and continuous tracks as seen in the top-down study (found in Figure 9). Both melt pool width and depth show similar trends and no key-hole welds were identified due to the lack of L.E.D applied. For each power setting, the largest widths and depths are seen during application of the lowest scan speed (i.e. highest energy density). Line instabilities and discontinuity become apparent as we see larger variations between each measurement taken in some parameter sets (e.g. 300W, 2.10 m/s). As L.E.D was decreased, widths and depths values reduced until a point at which no fusion took place (e.g. 300W, 2.10 to 3.00 m/s).1207(300W)(350W)(400W)0.30.60.91.21.5 Scan Speed (m/s)1.82.12.42.7Figure 10.

The generation of particles (as seen in Figure 14 (A)) is governed by the experimental PSD data applied. Particles generated ranged from 14.5 μm to 65.7 μm in diameter as seen the temperature legend (see Figure 14 (E)). Once generated, the particles fall and are numerically spread across the baseplate lowered a distance equal to one-layer thickness (i.e. 60 μm); in this instance the substrate used was a flat surface. Once the rigid spreader has surpassed the length of the table and all particles meet equilibrium, this forms the final powder bed geometry as seen in Figure 14 (F).(A)(B)(C)(D)Diameter - m6.572E-0056.5E-OO6E-0055.5E-0055E-0054.5E-005(E)4E-0053.5E-0053E-0052.5E-0052E-0051.5E-OO1.45E-005Figure 14. Particles are first generated (A) and then fall onto the spreading platform (B). After a pre-set time, the spread moves lineraly across the spreader and onto the build platform (C).

Particles are first generated (A) and then fall onto the spreading platform (B). After a pre-set time, the spread moves lineraly across the spreader and onto the build platform (C). The spread continues across the platform to completely disperse the powder (D). Once the powder has reached equilibiurm, this solid geometry was input into the melting model (E).1209Melting ModelESI-AM’s melting module produces a 3D-transient result that may be analysed at any time-step simulated. Seven cases were simulated with various parameter sets and model settings applied. Models 1 to 6 were deemed low-quality due to the course mesh applied to the geometry. Figure 15 shows some results from model 2; application of 400 W and 0.67 m/s, equating to a L.E.D of approx. 600 J/m.

Figure 15 shows some results from model 2; application of 400 W and 0.67 m/s, equating to a L.E.D of approx. 600 J/m. Within each 900 μm domain (for all models) the modulated laser was applied 9 times with a point distance of 70 μm to achieve a line length of 630 μm.When analyzing the results of models 1 to 6, it was realised that a miscalculation in Gaussian Beam radius resulted in the application of a spot size equal to almost half of that used in physical experimentation. Due to this, the resulting melt pools displayed a large amount of key-hole melting (as seen in Figure 15). In addition to this, the uncompensated point distance of 70 μm resulted in a huge amount of variation in melt pool depth as insufficient melting caused large gaps between welds (as seen in Figure 15 (B)). Throughout the laser application process a maximum temperature of 3466 K was reached in the centre of the melt pool as seen in Figure 15 (C).

Throughout the laser application process a maximum temperature of 3466 K was reached in the centre of the melt pool as seen in Figure 15 (C). Following such high temperatures and with the application of recoil pressure and evaporation models, spatter formation was observed as material was ejected from the melt pool in various directions as identified in Figure 15 (D). In addition, the reduced spot sized applied and consequent key-hole mode of melting further increased the rate of spatter formation from the melt pool.Due to the mistaken spot size applied in models 1-6, the resulting melt pool widths were much smaller for each parameter set applied which saw further deviation as the L.E.D was increased (as seen inFigure 17 (D)). In the physical experimentation, a maximum depth in excess of 300 μm was reached which surpassed the 200 μm substrate domain applied to the model.

In the physical experimentation, a maximum depth in excess of 300 μm was reached which surpassed the 200 μm substrate domain applied to the model. The excessive key hole melting led to depths beyond this maximum which could not be assessed thereby reducing the average depth. Heights measured in each model were a lot lower than that of physical experimentation. Again, without the right spot size and mesh quality applied, it is hard to judge as to why this may be. Although, this should be investigated further as the mesh quality may prove adequate in melt pool analysis with the correct spot size applied.Unfortunately, only a single high-quality model (model 7) was simulated due to time and computational constraints. Model 7 was simulated using the optimum parameters found for this alloy (as seen in Table 4). As well as the highly refined mesh, the application of a 70 μm spot size was applied by recalculating the Gaussian beam radius.

As well as the highly refined mesh, the application of a 70 μm spot size was applied by recalculating the Gaussian beam radius. By doing so, the melt pool width doubled to a value close to the experimental result. In addition, the variation of powder depth along the single line reduced dramatically to a similar variation to that found in physical experimentation. In this case the maximum temperature went beyond that which was estimated (3500 K) resulting in the lack of colour seen in the last weld of Figure 16 (C). Once again, both recoil pressure and evaporation model was applied resulting in the ejection of small particulates forming spatter Figure 16 (D). Both widths and depths measure in model 7 correlated well with the physical experimentation, but once again the height still seemed to be considerably smaller.

Both widths and depths measure in model 7 correlated well with the physical experimentation, but once again the height still seemed to be considerably smaller. Further studies must be conducted using a similar model setup (but various L.E.Ds applied) to further validate this model.1210900μm(B)T - K 346630002500(A)(D)(C)(E)(C)Figure 15. 3D transient results from model 2 (400 W, 105 μs): (A) top-down view of laser processing across powder bed, (B) side-view of melted cells, (C) temperature ditribution along melted track, (D) melted cells highlighting particle ejection due to material evaporation and (E) coloured legend indicating temperature range.900μm(B)T - K 3500(A)(D)(C)(E)Figure 16.

Due to the application of an inncorrect spot size (approx. half the actual value) to models 1-6, large deviation is seen between the models and experimental data. Model 7 (highlighted) included the correct spot size and concequently produced a realistic value.300250)m μ ( h t e p D2001501005000200400600800100012001400L.E.D (J/m)Models 1-6Exp.Model 7Poly. (Models 1-6)Poly. (Exp.)Figure 18. Melt pool width as a fuction of L.E.D for models 1 to 7 and experimental results. The maximum achievable melt pool depth was limited by the boundary applied in the melting model (200 μm). The combination of this and incorrect spot size led to a huge variation in depth in models 1 to 6. Model 7 (highlighted) included the correct spot size and concequently produced a similar value and variation to that found by physical experimentation.250200)m μ (t h g i e H15010050.. l ···· .·+·' ..... J ...........................

l ···· .·+·' ..... J ........................... , ············ •············• .... , ...................... .. !I!'···············y••············· t··············•!I!'00200400600800100012001400L.E.D (J/m)Models 1-6Exp.Model 7Poly. (Models 1-6)Poly. (Exp.)Figure 19. Melt pool height as a fuction of L.E.D for models 1 to 7 and experimental results.1212ConclusionIN625 was used to conduct a single line study in order to validate model results produced using ESI-AM’s spreading and melting modules. Process parameters were first optimised for the alloy by method of density measurement based on Archimedes principle. Using the optimised results, a solid rectangular geometry, known as ‘the crucible’, was manufactured to serve as a substrate in which the single lines were processed on. Three power settings were applied; 300 W, 350 W and 400 W.

Three power settings were applied; 300 W, 350 W and 400 W. For each power setting 10 exposure times were applied ranging between 23.33 μs and 233.33 μs (or 3.00 m/s to 0.30 m/s). This resulted on the application of 30 L.E.Ds varying from 100 J/m to 1333.33 J/m.Firstly, the line morphologies were studied to assess line quality. During the application of a high L.E.D, lines appear to be thick and stable. As the energy density is decreased instabilities result in track irregularities which caused balling. A further decrease resulted in little or no adherence due to insufficient power and/or exposure time to melt the powder.Secondly, melt pool morphologies were analysed along each line in 3 separate positions to obtain average width, depth and height measurements. Typically, application of a high L.E.D resulted in stable and continuous tracks as seen in the top-down study.

Typically, application of a high L.E.D resulted in stable and continuous tracks as seen in the top-down study. Both melt pool width and depth show similar trends and no key-hole welds were identified due to the lack of L.E.D applied. For each power setting, the largest widths and depths are seen during application of the lowest scan speed (i.e. highest energy density). Line instabilities and discontinuity become apparent as we see larger variations between each measurement taken in some parameter sets. As L.E.D was further decreased, widths and depths values reduced until a point at which no fusion took place.Using the spreading module, a powder bed was numerically generated using details relating to material properties, spreading geometry and particle size distribution. Once the spreading process had complete and powder bed had reached equilibrium, this final geometry was imported into the melting module.

Once the spreading process had complete and powder bed had reached equilibrium, this final geometry was imported into the melting module. Using the melting module, seven cases were simulated with various parameter sets and model settings applied. Models 1 to 6 were deemed low-quality due to the course mesh applied to the geometry and model 7 was deemed high-quality due to the refined mesh applied. Due to time and computation limitations, only one high quality model was simulated. In addition to the low quality of models 1-6, a miscalculation in Gaussian beam radius resulted in the application of a spot size half that applied in physical experimentation. Due to this, the results did not correlate well to that of the physical experimentation. Conversely, the single high-quality model simulated (model 7) showed promising results, although further simulations covering a range of L.E.Ds must be completed to fully validate the model.References[1]F.

COMPUTATIONAL STUDY SUMMARIZING THE EFFECTS OF COMPOSITION ON THE MELT POOL GEOMETRY IN ADDITIVE MANUFACTURINGDaniel Gwynn1, Sree Samhita Gundimeda2, Nandana Menon1, and Amrita Basak1, *1Department of Mechanical Engineering, The Pennsylvania State University, University Park, PA 16802, USA2 Green Valley School, Near Airforce Campus, Lekawada, Gujarat 382042, IndiaCommunicating Author: aub1526@psu.eduAbstractThermophysical properties play a crucial role in controlling the melt pool dimensions in metal additive manufacturing (AM). It is, therefore, imperative to understand the impact of thermophysical properties on the melt pool dimensions for critical materials such as nickel-based superalloys. In this paper, a three-dimensional analytical model is used to predict the steady-state melt pool dimension over a range of process parameters for several nickel-based superalloys.

In this paper, a three-dimensional analytical model is used to predict the steady-state melt pool dimension over a range of process parameters for several nickel-based superalloys. The effects of composition, in terms of the thermophysical properties, on the melt pool dimensions are also studied. The results show that the melt pool dimensions correlate well with the liquidus temperature, density, and thermal conductivity of the alloy. By exploring the impact of process parameters and compositions on the melt pool dimension evolution, a framework can be established to maintain the desired melt pool dimensions during the fabrication of functionally- graded parts with different alloys belonging to the same material class.IntroductionNickel-based superalloys are widely used in the production of gas-turbine hot-section as they provide higher yield strength with increasing operating temperatures partly due to the presence of secondary γ′ phases [1].

The hot-section components are manufactured using conventional manufacturing methods such as investment casting. However, the fabrication of complex designs and functionally-graded components is not feasible using such conventional techniques. Additionally, repair of the superalloy components using conventional methods is challenging as they are difficult to weld. By addressing both these constraints, metal additive manufacturing (AM) presents itself as a viable solution.Metal AM works on the principle of consolidating metallic feedstocks (e.g., wire or powder) in layers using a high-energy heat source (e.g., a laser, an electron beam, or an arc). Understanding the fundamental physics of metal AM is a critical requirement for process development, optimization, and control [2]. The heat source in metal AM (e.g., powder bed fusion, directed energy deposition) creates a small melt pool that controls the microstructures and, therefore, the mechanical properties of the final part [3].

This makes melt pool geometry a critical parameter in metal AM. Numerous studies involving the modeling and simulation of the melt pool geometry in metal AM processes for nickel-based alloys are available in literature. Analytical778models are extensively used to model the melt pool geometry for nickel-based superalloys. Gäumann [4] used a modified Rosenthal solution [5] to obtain the liquid isotherm and a circular melt pool. Christensen et al. [6] suggested an elliptic melt pool along the direction of the beam’s movement. Goldak et al. [7] assumed a double-ellipsoid weld pool shape that consisted of front and rear ellipsoid quadrants. Wang et al. [8] developed a finite element model expanded to include an intensity distribution for estimating the melt pool dimensions of CMSX-4®. More recently, Mondal et al.

Wang et al. [8] developed a finite element model expanded to include an intensity distribution for estimating the melt pool dimensions of CMSX-4®. More recently, Mondal et al. [9] developed a Gaussian Process-based regression model to predict the melt pool geometry for a popular nickel-based superalloy, CMSX-4®.In addition to analytical models, Finite Element Methods (FEM) and Finite Volume Methods (FVM) are also used for predicting the melt pool geometry in metal AM of nickel-based superalloys [10]. A computational fluid dynamics (CFD)-based model allows the incorporation of fluid convection inside the melt pool. To accurately simulate the condition inside the melt pool, surface tension-driven flow known as the Marangoni convection (or Bernard-Marangoni convection) needs to be included in the flow-thermal model [3]. A liquid with a higher surface tension pulls more strongly on the surrounding liquid than the one with lower surface tension.

A liquid with a higher surface tension pulls more strongly on the surrounding liquid than the one with lower surface tension. This extends the melt pool laterally and makes it shallower. Acharya et al. [3] developed a three- dimensional CFD model and obtained good agreement with experimental data for the melt depth in CMSX-4® processed through powder bed fusion. A 2D approach was also implemented to simulate the melt pool characteristics for IN100 and obtained a reasonably good agreement for the melt depth and the average melt temperature [11]. This model was also extended by Basak et al.

This model was also extended by Basak et al. [12] to simulate melt pool characteristics for CMSX-4® with reasonable congruence between the simulation and the experimental results achieved for the melt depth and the average melt pool temperature.While several studies outline the effects of processing parameters on the melt pool geometry, there exists a knowledge gap in the exploration of the combined effects of the composition of nickel-based superalloys and process parameters on the melt pool geometry. This paper attempts to evaluate the effects of both composition and process parameters on the melt pool geometry of several different nickel-based superalloys. An analytical model presented by Eagar- Tsai [13] is utilized for predicting the melt pool geometry. The model is thoroughly validated using the literature data. Thereafter, several popular nickel-based superalloys are selected, and their thermo-physical properties are evaluated using JMatPro® [14].

The model is thoroughly validated using the literature data. Thereafter, several popular nickel-based superalloys are selected, and their thermo-physical properties are evaluated using JMatPro® [14]. The Eagar-Tsai analytical model is, thereafter, interrogated to evaluate the effects of composition and process parameters on the steady-state melt pool geometry (e.g., depth). The results show that the melt pool dimensions correlate well with the liquidus temperature, density, and thermal conductivity of the alloy. By exploring the impact of process parameters and compositions on the melt pool dimension evolution, a framework can be established to maintain the desired melt pool dimensions during the fabrication of functionally-graded parts with different alloys belonging to the same material class.MethodsThis section, first, provides a brief description of the analytical model.

Thereafter, themodel is validated against two sets of experimental results.779Analytical ModelThe analytical model selected in this work was originally developed by Eagar and Tsai [13]. This model is based on Rosenthal’s steady-state model. However, the model also uses a more realistic Gaussian beam power source instead of Rosenthal’s point source. Except for a point source and steady-state conditions, Rosenthal’s original assumptions remain intact [5]. Eagar- Tsai’s model starts with the following equation:𝑣𝑣(𝑤𝑤+𝑅𝑅) 2𝑎𝑎−𝑞𝑞 2𝜋𝜋𝜋𝜋𝜋𝜋𝑇𝑇 − 𝑇𝑇0 = −𝑒𝑒Here,is the input power (W), (m) is the distance from the heat source defined as 𝑞𝑞 - is the thermal diffusivityis the initial temperature (°C),is the temperature (°C),is the thermal conductivity (W/(m°C)), 𝑇𝑇0𝑇𝑇is the distance (m) in the,is the power source travel speed (m/s), 𝜋𝜋𝜋𝜋 , and direction in a moving coordinate of speed (𝑤𝑤 𝑤𝑤 𝑣𝑣 (m2/s).

(A) Schematic of the coordinate system used in the analytical model and (B) Definition of the melt pool depth, length, and width.780(1)𝑥𝑥(2)This model’s primary assumptions include the following:Convective and radiative heat transfer mechanisms are ignored. • The thermal properties of the material are constant. • The computational domain is semi-infinite. • The phase change of the material is not included.Despite its limitations, Eagar-Tsai’s model is extensively used for parameter space exploration through the design-of-experiments (DoE)-based approaches [9]. Computationally efficient surrogates have also been developed using this model for AM applications [9].Validation of the Analytical ModelThis section summarizes the validation results obtained using two different sources of experimental data. In the first validation, the melt pool measurements reported by Ning et al. [15] for Inconel 625 are utilized.

In the first validation, the melt pool measurements reported by Ning et al. [15] for Inconel 625 are utilized. Three different process parameter combinations are selected: Case 1 = 800 with power, is mm/s, and Case 3 with is calculated as = 0.4, an average error of 4.3% with a maximum error the laser absorptivity coefficient. Utilizing 𝑣𝑣 of 7.3%, and a minimum error of 1.5%, are observed for the melt pool length and depth when compared with the results reported by Ning et al. The results of this validation are shown in Figure 2.= 169 W and laser scan speed,= 875 mm/s; Case 2 with= 182 W and ×= 195 W and= 800 mm/s. Here,where 𝑣𝑣𝑃𝑃𝑣𝑣𝑃𝑃𝑃𝑃𝑃𝑃𝜂𝜂𝑞𝑞𝜂𝜂t e g n N m o r fi5045x=yg n N m o r fi)400375x=y)m µ ( h t p e Dl o o P t l eM. l a403530Case 1Case 3Case 2m µ ( h t g n e Ll o o P t l eM.

l a403530Case 1Case 3Case 2m µ ( h t g n e Ll o o P t l eM. l a t e350325300Case 1Case 3Case 23040 Melt Pool Depth (µm) from Current Work (A)354550300400 350 Melt Pool Length (µm) from Current Work325375(B)Figure 2. Validation results (A) for melt pool depth and (B) melt pool length using measurements from Ning et al. [14].For the second validation, the melt pool measurements recorded for CMSX-4® by = 1 mm/s; Case = 360 W andAnderson et al. [10] are utilized. Results are obtained for Case 1: 2: = 360 W and = 6 mm/s. The results of this validation are illustrated in Figure 3. This validation results in a 𝑣𝑣 maximum error of 21.5%, and a minimum error of 1.0%, cumulating to an average error of 9.4%.= 240 W and = 1 mm/s; and Case 4: 𝑣𝑣 𝑃𝑃= 240 W and= 6 mm/s; Case 3:𝑃𝑃𝑃𝑃𝑣𝑣𝑃𝑃𝑣𝑣781𝜂𝜂m o r f1000x=ym o r f2500)m µ ( h t [ e Dl o o P t l eM.

These properties are recorded in Table 2.Table 2. Material properties of composition of ten alloys selected in Table 1. Here ρ in (kg/m3),𝜌𝜌, 𝑘𝑘, 𝑎𝑎,𝑐𝑐 a in (m2/s10-6), k in (W/(mC), c in (J/(kgC), TL in (C)ρ aB1900 CMSX-4®8720 3.148592 3.24× GTD 4448488 3.12° IN 100 Mar M 200 + Hf 8559 3.337871 3.15° Mar M 0028563 3.32° PWA 14268487 3.41PWA 14848770 3.28Rene N68926 3.44k c11.97 436.7811.90 426.6211.09 418.6011.34 455.9511.97 419.4311.94 419.0112.54 432.0611.99 416.2612.45 405.52TL 135413791354135113761356138413791401To highlight the effect of process parameters, three different combinations of laser power, , are studied for a representative alloy, PWA-1426. The cases presented = 450 = 6 mm/s.

Additionally, amongst the three melt pool dimensions, the melt pool length appears to be the most influenced by a change in process𝑃𝑃𝑣𝑣𝑃𝑃𝑣𝑣783parameters. With this preliminary understanding of the impact of process parameters, the melt pool dimensions are analyzed over the range of alloys differing in composition, at constant . Figure 5 shows the depth variation of all ten alloys for Case B: = 6 mm/s. The 𝑃𝑃 𝑣𝑣 results show that for the same processing parameters, GTD 444 has the largest melt pool dimensions while SR-99 has the smallest with the difference in the melt pool depth, length, and width being 6%, 25%, and 12%, respectively.= 900 W and𝑃𝑃𝑣𝑣and)m m( n o i s n e m Di1086Depth Length Widthl o o P t l e42M0Figure 5.

Variation of melt pool dimensions for ten alloys processed at Case B: = 6 mm/s.= 900 W and𝑣𝑣 To understand the differences in the melt pool dimension with composition, the melt pool dimension (depth) is studied as a response to the thermo-physical properties of all ten alloys as investigated in Figure 5. These results are consolidated in Figure 6. Figure 6(A) shows the variation of the melt pool depth with the liquidus temperature, . The results show a significant negative correlation with . The temperature is the highest near the laser source and decreases away from the laser source. Since the definition of the melt pool boundary is the iso-surface of the liquidus temperature, a lower liquidus temperature would imply that the melt pool boundary would lie further away from the laser source, thereby, increasing the melt pool depth. Figures 6(B) and (C) illustrate the variation of melt pool depth with density, , respectively.

Figures 6(B) and (C) illustrate the variation of melt pool depth with density, , respectively. For an identical laser power, the material with a lower density would have a higher volumetric heat flux which would, in turn, increase the melt pool depth. The effects of the thermal conductivity are observed to have the most prominent influence on the ten alloys tested. With an increase in thermal conductivity, the melt pool gets elongated, due to an improved heat flow via conduction. This reduces the penetration and hence the melt pool depth is smaller with an increase in thermal conductivity.𝑃𝑃𝑇𝑇𝐿𝐿𝑇𝑇𝐿𝐿, and thermal conductivity,𝜌𝜌, and the thermal diffusivity, , on the melt pool depth as shown in Figures 6(D) and (E). However, no correlations are observed between the specific heat and the thermal diffusivity.

However, no correlations are observed between the specific heat and the thermal diffusivity. One would expect the influence of thermal diffusivity to be prominent due to an improved heat flow via diffusion which would increase the area of the melt pool along the scan direction and, hence, decrease the melt pool depth. The specific heat controls the rate at which the material would respond to heat flux therebySimilar plots are also made to investigate the effects of the specific heat,𝑐𝑐𝑎𝑎784𝑘𝑘suggesting a lower specific heat may contribute to an increase in the melt pool size.

Models of higher fidelity are, thus, required to capture the relationship between the thermophysical properties and melt pool with better accuracy and reliability.)m m( h t p e Dl o o P t l eM1.651.601.551.50y = -0.0009x + 2.857 R² = 0.68581350138014101440)m m( h t p e Dl o o P t l eM1.651.601.551.50y = -5E-05x + 2.0343 R² = 0.3999700080009000 10000TL (°C) (A)(B)ρ (kg/m3))m m1.65( h t p e Dl o o P t l eM1.601.551.5010y = -0.0321x + 1.9905 R² = 0.545311121314k (W/m°C) (C)(D))1.65m m( h t p e Dl o o P t l e1.601.551.50y = -0.0789x + 1.8622 R² = 0.1029M3.003.203.403.60a (m2/s) ⅹ 10-6 (E)Figure 6. Variation of melt pool depth with (A) fit and R2 values are also indicated on each plot. 𝑇𝑇𝐿𝐿, (B)𝜌𝜌, (C)𝑘𝑘, (D)𝑐𝑐, and (E)𝑎𝑎. The dotted linear) for also confirm the significance of predictions.

𝑇𝑇𝐿𝐿, (B)𝜌𝜌, (C)𝑘𝑘, (D)𝑐𝑐, and (E)𝑎𝑎. The dotted linear) for also confirm the significance of predictions. However, both specific heat and thermal diffusivity do not appear to be significant ( ) for the ranges evaluated. This gives an important insight into what variables are critical. E.g., While building predictive models for melt pool geometry, specific heat and thermal diffusivity could be avoided. Including non- significant variables can affect the accuracy and increase the dimensionality of the input space, thereby requiring more data.Thevalues for the regressions are also reported in Table 3.

The smallvalues (,and 𝑝𝑝𝑝𝑝𝑇𝑇𝐿𝐿𝜌𝜌𝑘𝑘𝑝𝑝 > .05< .05785Table 3.values reported for the linear regression.Term𝑝𝑝 CoefficientVariableVariable TL ρ k c a0 0 0 .001 0.001 .05 .015 .923 .366ConclusionA three-dimensional analytical melt pool model based on the work of Eagar-Tsai is developed and validated with results reported by Ning et al. and Anderson et al. with predictions resulting in average errors of 4.3% and 9.4%, respectively. These results are obtained after appropriate calibration of the beam distribution parameter and/or the laser absorptivity coefficient. The model is interrogated to evaluate the effects of the thermophysical properties resulting from the different alloy compositions on the melt pool dimension. The melt pool dimensions show significant correlations with density, liquidus temperature, and thermal conductivity.

The melt pool dimensions show significant correlations with density, liquidus temperature, and thermal conductivity. It is anticipated that this model can provide the desired fidelity for quick screening of alloy compositions and process parameters.While the developed model can provide sufficiently accurate results, its accuracy may be improved by including additional parameters at the cost of added complexity. Some of these additions that would be expected to improve model accuracy include using temperature-based thermophysical material properties, including provisions for finite dimensions, incorporating convective and radiative heat transfer, and adding latent heat effects due to phase change. In the future, surrogate models [9] will be developed over a wide range of alloy compositions. Since the alloy composition is a high-dimensional space (i.e., the number of elements in an alloy is fifteen or more), suitable dimensional reduction methodologies need to be implemented.

Since the alloy composition is a high-dimensional space (i.e., the number of elements in an alloy is fifteen or more), suitable dimensional reduction methodologies need to be implemented. The Eagar-Tsai model is not accurate. To improve the accuracy of the prediction, multi-fidelity surrogates will also be developed [17].AcknowledgmentsThe work reported in this paper is supported by the Department of Mechanical Engineering at the Pennsylvania State University, University Park, PA 16802. Nandana Menon would like to acknowledge the James E. Marley Fellowship, the Kulakowski Travel Grant awarded by the Mechanical Engineering Department of the Pennsylvania State University, and the National Science Foundation Student Support offered by the organizing committee of the Solid Freeform Fabrication Symposium. Any opinions, findings, and conclusions in this paper are those of the authors and do not necessarily reflect the views of the supporting institutions.786References[1] N. M.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCOMPUTED AXIAL LITHOGRAPHY FOR RAPID VOLUMETRIC 3D ADDITIVE MANUFACTURINGBrett	E.	Kelly1,3,	Indrasen	Bhattacharya2,	Maxim	Shusteff3,	Hayden	K.	Taylor1,	and	 Christopher	M.	Spaddacini31Department	of	Mechanical	Engineering,	University	of	California,	Berkeley.	Berkeley,	 2Department	of	Applied	Science	and	Technology,	University	of	California,	Berkeley	 3Materials	Engineering	Division,	Lawrence	Livermore	National	LaboratoryAbstractThe vast majority of additive manufacturing processes today operate by printing voxels serially point-by-point to build up a 3D part. In some higher throughput techniques, for example optical printing methods such as projection stereolithography [1], [2], parts are printed layer-by-layer by curing full 2D (very thin in one dimension) layers of the 3D part in each print step.

In this work, we demonstrate a new technique which prints entire parts all at once and eliminates layering. The approach, termed Computed Axial Lithography (CAL), is based on tomographic reconstruction, with mathematical optimization to generate a set of projections to optically define an arbitrary dose distribution within a target volume and to cure the entire volume simultaneously. Volume-at-once fabrication of complex geometries is achieved using a custom system built from a DLP projector and a rotating resin volume. This technique can be used to expand the range of printable geometries in additive manufacturing and relax constraints on factors such as overhangs in topology optimization. The method also vastly increases print speed for 3D parts. We show complex and overhanging geometries printed in situ without layering.IntroductionAdditive manufacturing (AM) technology has expanded greatly since the inception of stereolithography in the 1980s.

Today, AM techniques are used in a variety of fields to print structures that often could not be manufactured by conventional subtractive processes. Applications range from biological scaffolds for culturing tissue in vitro[3] to micro-architected materials with microstructures designed to produce enhanced or novel material properties [4]. Most AM methods today operate in one of a series of forms which we can categorize by the geometric complexity of their unit print operation, or the geometry of the material that is printed in a single print step. A majority of printing methods today rely on a print step which is either zero or one-dimensional. Zero- dimensional (0D) print operations include those where material is formed or processed at a single point within the print volume. Methods such as selective laser sintering (SLS), selective laser melting (SLM), stereolithography (SLA) and inkjet printing fall into this category.

Methods such as selective laser sintering (SLS), selective laser melting (SLM), stereolithography (SLA) and inkjet printing fall into this category. Another large subset of methods exist which rely on a unit print operation which is one dimensional (1D), or which print a line in the unit operation. These techniques include, most prominently, extrusion-based methods such as fused deposition938modeling (FDM) and direct ink writing (DIW). Recently, more and more techniques are emerging which use a two-dimensional (2D) unit print operation. These techniques are mostly optical and include projection stereolithography methods [1] and Continuous Liquid Interface Production (CLIP) [2]. The aforementioned 0D, 1D, and 2D unit operation printing methods generally achieve printing of 3D geometries by repeating their unit print operations serially in the remaining dimension(s).

The aforementioned 0D, 1D, and 2D unit operation printing methods generally achieve printing of 3D geometries by repeating their unit print operations serially in the remaining dimension(s). In the 2D case, this involves repeating the operation once for sequential thin cross-sections or “layers” of the part. Recently, the authors of this work introduced a technique capable of printing entire 3D parts in a single step [5]. This technique was based on the use of intensity superposition of beams from three orthogonal directions to photopolymerize a 3D geometry in situ in a single print step. It introduced a new paradigm of volumetric fabrication which provides a number of potential advantages over layer-by-layer printing. However, the three beam superposition technique, while a useful demonstration of the ability to print volumes at once, is constrained in the range of 3D geometries it can print to those that can be formed by intersecting three orthogonal images.

This paper introduces a new technique, Computed Axial Lithography (CAL), which preserves the ability to form entire parts at once in situ, but greatly expands the geometric flexibility.The ability to print full 3D parts volumetrically, with a three-dimensional unit print operation, offers a number of advantages over layered printing. To start, the CAL technique has already demonstrated the ability to print 3D parts faster than layered methods. In a 15mm x 15mm x 15mm print volume, the complex parts in Figures 8A and 8C were printed in under 8 minutes while the sphere geometry in Figure 8B was printed in just 80 seconds. More interestingly, volumetric printing offers unique potential advantages in the ability to print geometries that cannot be printed layer-by-layer. These geometries include those which require support material in layered printing, such as geometries with overhangs and the geometry demonstrated in Figure 8A.

These geometries include those which require support material in layered printing, such as geometries with overhangs and the geometry demonstrated in Figure 8A. Finally, by printing parts all at once, rather than in sequential layers, mechanical anisotropies arising from layering can be reduced or eliminated.The CAL approach described in this work is founded on the generation of an accumulated 3D intensity distribution designed carefully in conjunction with photocuring chemistry such that in a single unit operation with a single development step, a 3D part of arbitrary user-defined geometry is generated. It is rooted in the basic concepts of computed tomographic (CT) imaging [6], which we apply to stereolithography.CAL Printing AlgorithmRather than printing the target 3D geometry layer-by-layer, the CAL method is designed to expose the transparent resin by projecting 2D images from an array of angles about an azimuthal axis.

A major component of this research was the development of a computational algorithm to design the projected image from each angle based on a target 3D geometry. The optical design method for CAL is based on one of the possible reconstruction procedures in computed tomography (CT). Prior art in a similar technique exists in intensity modulated radiation therapy (IMRT) for cancer treatment [7]. IMRT seeks to produce a 3D variation of radiation dose in a target volume in the patient's body. Typical practice in this radiotherapy procedure involves the projection of a small number (7-10) of 2D distributions of X-ray pencil beams at a few selected angles. The intensity of939the beamlets is chosen to satisfy a set of therapeutic constraints including: sufficient but not excessive dosage in the tumor, low radiation dose in certain critical organs as well as the physical constraint that radiation dose must be positive.

This is a challenging inverse problem, for which iterative optimization techniques have yielded success in defining previously impossible dosage distributions, including concave regions such as the prostate gland. Producing dose distributions in a clinically reasonable timeframe while satisfying multiple objectives continues to be an area of active research. From a mathematical perspective, the volumetric 3D printing problem has similarities in terms of constraints, but greater freedom in the exact dose distribution because of the non-linear thresholding behavior of the photoresins used.To describe the optical design algorithm, it is easiest to start by reducing the problem of generating 2D projections to print a 3D geometry to that of generating 1D projections to print a 2D geometry. Consider the 2D geometry to be a slice of the full 3D part.

Consider the 2D geometry to be a slice of the full 3D part. This 2D flatland analysis is then readily extended to the third dimension by concatenating calculated 1D projections for subsequent slices of the 3D part. Note that this technique is still distinctly different from layered printing as neighboring layers form simultaneously rather than sequentially. In the raytracing picture used in this algorithm, we conisder all the beams for a particular projection angle to be parallel. This is analogous to parallel beam tomography in the medical community.

In the raytracing picture used in this algorithm, we conisder all the beams for a particular projection angle to be parallel. This is analogous to parallel beam tomography in the medical community. In the CT imaging configuration where a uniform pencil beam at azimuthal angle 𝜃 is projected into the 2D imaging volume with optical density given by 𝑅 𝑥, 𝑦 , the collected dose distribution on the 1D camera along the space dimension x is given by 𝑃(𝑥, 𝜃):𝑃 𝑥, 𝜃 = 𝑅 𝑥 cos 𝜃 − 𝑢 sin 𝜃, 𝑥 sin 𝜃 + 𝑢 cos 𝜃 𝑑𝑢where −𝑢 sin 𝜃, 𝑢 cos 𝜃 for parameter 𝑢 represents a line through the origin in the direction of exposure for the particular angle 𝜃 . 𝑃(𝑥, 𝜃) is the well known Radon transform of the 2D image 𝑅. We will refer to 𝑃(𝑥, 𝜃) as the angular projections of the image. From the projection slice theorem, it turns out that the Fourier transform of a projection at angle 𝜃 is exactly equal to a 1D sample of the original image's 2D Fourier transform 𝑅(𝑘!, 𝑘!).

From the projection slice theorem, it turns out that the Fourier transform of a projection at angle 𝜃 is exactly equal to a 1D sample of the original image's 2D Fourier transform 𝑅(𝑘!, 𝑘!). The 1D sample is taken along a slice 𝑘 cos 𝜃 , 𝑘 sin 𝜃 corresponding to the particular angle at which the projection was integrated. This can be expressed as:𝑃 𝑥, 𝜃 𝑒!!"#𝑑𝑥 = 𝑅 𝑘 cos 𝜃 , 𝑘 sin 𝜃This is illustrated in Figure 1, where 1D cuts in the projection space image can be compared to similarly shaded central slices in the Fourier domain. Sufficiently dense sampling in the Fourier domain is required for an accurate CT image reconstruction.940(1)(2)Figure	1	–	Representation	of	Fourier	slicing	and	sampling	based	on	angular	projections	from	many	 anglesThe reconstruction of the 2D image volume can be understood as an algorithmic time reversal of the CT imaging process.

This back-projection algorithm starts with the measured 1D projections and propagates each individual projection backwards while uniformly exposing the target region with this intensity pattern. This is repeated for every angle. From the central slice theorem, this corresponds to building up the sample slice by slice in the Fourier domain. This algorithmic back-projection then motivates a technique to physically back-project the computed Radon transform at each angle and directly construct desired 3D dose volumes. However, directly back-projecting the 2D radon transform will not produce the correct result. In order to compensate for the inverse radial oversampling inherent in the Fourier slicing approach, we apply a radially increasing ramp filter. We also apply a window to the ramp filter in the Fourier domain, so as to exclude high spatial frequencies beyond the degree of sampling provided by the number of angular samples.

We also apply a window to the ramp filter in the Fourier domain, so as to exclude high spatial frequencies beyond the degree of sampling provided by the number of angular samples. IMRT literature suggests an exponential windowing filter for a smooth backprojection filter. This approach leads to the following backprojection filter in the Fourier domain (𝑘, 𝜃):𝐻 𝑘 = 𝑘 𝑒!! !!!941(3)where 𝑘! is chosen based on the number of angular samples. This windowed high pass filter is applied on every Fourier slice before back-propagating it to form the image. The spatial domain representation of the filter has negative ripples that often lead to the backprojections being negative even if the projections themselves are positive. This is a challenge since the backprojections are physically constrained to positive values.

This is a challenge since the backprojections are physically constrained to positive values. We address this initially in two ways: 1) by adding a constant offset to all projections to raise the minimum value to zero, or 2) by setting all negative projection values to zero. We then use this set of projections as an initialization step for a constrained optimization procedure.The goal of the optimization algorithm is to calculate the set of backprojections 𝑃!"# 𝑥, 𝜃 that best produces a desired output intensity. We use an iterative optimization procedure based on projected gradient descent. This method is guaranteed to converge for a convex objective and convex constraint set. Neither of these is true in our case due to the thresholding property of the resin and discrete values for the projector input. However, this heuristic performs quite well for some simple geometries, and reasonably well for more complicated ones.

However, this heuristic performs quite well for some simple geometries, and reasonably well for more complicated ones. One iteration of the optimization loop to generate 𝑃!!! 𝑥, 𝜃 from 𝑃! 𝑥, 𝜃 given a target image 𝑅(𝑥, 𝑦) goes as follows (illustrated in Figure 2):Figure 2 - The optimization flow, with example images corresponding to each stage in the algorithm.1) Projection: Starting with the 8-bit DLP projections 𝑃! 𝑥, 𝜃 computationally generate an unthresholded 2D dose distribution of power 𝐷(𝑥, 𝑦). This requires an accurate understanding and calibration of the system’s forward model. The relationship between specified pixel brightness values and the resulting projected intensity is extracted from measurements taken with a calibrated silicon photodiode.

The relationship between specified pixel brightness values and the resulting projected intensity is extracted from measurements taken with a calibrated silicon photodiode. We assume, for now, a942chemical model where regions which receive more than a critical threshold dose form solid, crosslinked material while those with less than the critical dose remain uncrosslinked and are later developed away. A future implementation will directly solve photopolymerization rate equations integrated over time and space to generate the degree of crosslinking in the forward model.2) Thresholding: Depending on the development recipe, convert the degree of crosslinking to the thresholded image at the nth iteration 𝑅! 𝑥, 𝑦 . W have experimented with variations of this procedure, to prevent the optimization from being sensitive to a single threshold, and instead have attempted to penalize absolute errors around the threshold as well.

In all cases however, 𝑅!(𝑥, 𝑦) lies between 0 and 1, representing material fully washed away as opposed to fully present3) Error: We determine the error in the image by comparing with the target:𝛿! 𝑥, 𝑦 = 𝑅! 𝑥, 𝑦 − 𝑅(𝑥, 𝑦).We transform this into the backprojection domain by performing a Radon transform (integral projection) at every angle, followed by ramp filtering. This leads to the projection domain error 𝛿! !(𝑥, 𝜃)4) Update: An unconstrained new set of projections is computed as: 𝑅!!! 𝑥, 𝜃 = 𝑅!(𝑥, 𝜃) − 𝛿! !(𝑥, 𝜃). Finally, the computed projection is constrained to positive 8-bit values by first setting negatives to zero, and then quantizing. It can be confirmed that the updated 𝑅!!!(𝑥, 𝜃) is the closest element in the constrained set to the computed unconstrained value and is therefore a projection Algorithm ValidationAs an incremental step before applying the optical design algorithm toward 3D CAL printing, a physical validation of the algorithm was performed.

For this validation, anAAlgorithm Validation SystemBExample: image frames from projector angle θjangle θiCExample:Sum of all movie framesThin layer of resinLens3D slice analogs for each video frame3D analog: net exposure dose through for slice over one rotationDLP ProjectorFigure	3	-	A:	Physical	system	used	to	validate	optical	design	algorithm.	B:	Example	movie	frames	 from	projected	video	and	representation	of	analogy	to	3D	system.	Each	frame	corresponds	to	the	 dose	passing	through	the	resin	from	a	particular	angle	𝜽.	C:	Example	sum	of	movie	frames	from	all	 angles.	This	dosage	distribution	is	analogous	to	the	net	exposure	dose	seen	by	a	horizontal	slice	of	 the	3D	volume	over	one	rotation.943experimental apparatus was constructed as depicted in Figure 3A. In this apparatus, a DLP projector was used as the optical source to generate the projections. The system was configured such that a 2D image was incident on a thin layer of resin.

In this apparatus, a DLP projector was used as the optical source to generate the projections. The system was configured such that a 2D image was incident on a thin layer of resin. The video output from the projector was set to match the time-evolving intensity map which would be incident upon a cross-section of the 3D volume in the 3D CAL system. Thus, each video frame corresponded to a 1D projection propagated in a single direction dictated by angle of illumination from the algorithm as depicted in Figure 3B. As designed by the algorithm, the sum of all video frames gives an image which approximates the target geometry. An example is shown in Figure 3C.The algorithm was tested for a range of geometries. Projections from 500 evenly spaced angles about 360o were computed and used to generate the video frames.

An example is shown in Figure 3C.The algorithm was tested for a range of geometries. Projections from 500 evenly spaced angles about 360o were computed and used to generate the video frames. The frame rate of the projected movie was set to simulate an angular rotation of 25o/s to match the max rotational speed of the rotating stage used later in the 3D CAL printing system. This setup yielded useful results to validate the algorithm. Example print results are shown in Figure 4.Figure	4	-	Algorithm	validation	results.	For	four	geometries,	the	summed	dose	from	all	angles	and	 the	corresponding	print	results	are	shown	A-B:	a	geometry	with	sharp	edges.	C-D:	a	geometry	with	a	 void	inside	of	a	solid	object.	E-F:	a	geometry	with	a	solid	piece	inside	another	solid	piece.	G-H:	a	 complex	geometry	with	curved	features.Extension to 3DIn order to translate the optical design algorithm from 2D validation to 3D printing, additional physical effects must be considered.

These effects arise from the fact that in 3D, the projections from each angle propagate through the resin volume and encounter attenuation and changes in refractive index. The algorithm does not currently account for these effects, although it soon will. Rather than accounting for these effects in the algorithm, we have addressed them in the design of the photoresin. First, rather than modelling attenuation in the algorithm, we sought to minimize it. This was achieved by carefully selecting the photoinitiator concentration used in the resin. Based on the photoinitiator concentration in the resin, a tradeoff exists between intensity uniformity and print speed. A low photoinitiator concentration minimizes absorption and preserves944intensity uniformity of the projection through the volume, but also leads to slower print times. The slow print time, however, can be counteracted by using higher intensity.

The slow print time, however, can be counteracted by using higher intensity. In the prints shown two photoinitiators were used: Irgacure 819 and Camphorquinone (CQ). The absorption spectra of each of these constituents were measured on a UV-Vis spectrophotometer and are plotted in Figure 5, along with the spectral output of the blue channel of the DLP projector used.inthis work,To address the change in refractive index, we rely on the geometry being defined during the oxygen inhibition phase of polymerization [8],[9]. During the oxygen inhibition phase, generated photoradicals quench local oxygen molecules faster in regions of higher dose. At the onset of curing, when the optical properties begin to change, the geometry has already been defined through spatially localized depletion of oxygen. Since a large fraction of the optical dose is used to define local regions that are free of oxygen, we may say that much of the 3D printing process is invisible to the eye.

Since a large fraction of the optical dose is used to define local regions that are free of oxygen, we may say that much of the 3D printing process is invisible to the eye. This has the advantage that parts that have been ‘printed’ already do not lead to significant optical scattering as the process continues.Figure	5	–	Absorption	spectra	of	the	two	photoinitiator	molecules	used	in	this	work	plotted	against	 the	spectral	output	of	the	DLP	projector	used.The final effects that arise in the 3D printing case but not in the 2D algorithm validation case are the emergence of gravitational, buoyant, and centripetal forces as the curing part increases in density compared to the uncured resin. To minimize the effect of forces exerted on the curing part, a high viscosity resin was formulated. The polymer used in this resin comprised of a mix of 75 wt% Bisphenol A glycerolate (1 glycerol/phenol) diacrylate and 25 wt% PEGDA 250 Da.

The polymer used in this resin comprised of a mix of 75 wt% Bisphenol A glycerolate (1 glycerol/phenol) diacrylate and 25 wt% PEGDA 250 Da. The viscosity was measured using cone and plate rheometry to be ~4000 cP and independent of shear rate. All polymer constituents and photoinitators were obtained from Sigma Aldrich.945CAL 3D PrintingIn order to demonstrate the ability of CAL to print 3D parts, a 3D printing system was constructed. This system is depicted in Figure 6A. The same DLP projector was used as the optical source to generate the projections. A cylindrical glass vial was used to contain the resin. This vial was fixed from the top to a controlled rotating stage. In the printing system, the image output of the projector is synchronized to the rotation of the vial.

This vial was fixed from the top to a controlled rotating stage. In the printing system, the image output of the projector is synchronized to the rotation of the vial. The image projected at a particular angle contains a concatenation of 1D projections for each vertical slice of the target geometry to form a 2D image as shown in Figure 6B, where each frame is the image output from the algorithm for a particular angle. To avoid cylindrical lensing effects which distort the incident image, a rectangular box filled with a fluid of refractive index > 1 was used.

To avoid cylindrical lensing effects which distort the incident image, a rectangular box filled with a fluid of refractive index > 1 was used. In this work, the index-matching fluid used was the same prepolymer composition used in the resin but without the photoinitiator added.Figure	6	-	CAL	3D	printing	system.	A:	full	system	schematic.	B:	an	example	of	an	image	frame	output	 from	the	projector	for	a	particular	angle.	C:	an	image	of	a	3D	structure	forming	inside	the	system.946Using the 3D CAL printing system, successful 3D prints of various geometries were achieved. As a first example, parts with uniform vertical cross-section were printed. In these parts, the algorithmic design of the projections for each vertical slice is the same, making the computation simpler and faster. These results are shown in Figure 7. The first geometry, a rectangular prism, demonstrates the potential of CAL to print 3D geometries with sharp corners.

These results are shown in Figure 7. The first geometry, a rectangular prism, demonstrates the potential of CAL to print 3D geometries with sharp corners. Another geometry, with a semicircle cross-section, demonstrates the ability to simultaneously print curved surfaces, flat surfaces, and corners.

Another geometry, with a semicircle cross-section, demonstrates the ability to simultaneously print curved surfaces, flat surfaces, and corners. Most importantly, successful printing of a hollow cylinder demonstrates the ability to print voids within a solid structure.Figure	7	-		3D	CAL-printed	results	with	constant	cross-section	in	z.	Column	1:	1D	projections	vs.	 angle	for	each	z	slice.	Column	2:	summed	intensity	distribution	from	all	projections	in	each	z	slice.	 Column	3.	Images	of	3D	printed	results947Figure	8	-	Examples	of	3D	printed	parts	with	non-uniform	cross-section.	Algorithm	output	is	 shown	as	a	point	cloud,	plotting	points	where	the	net	exposure	dose	exceeds	a	threshold.	Color	 represents	the	axial	(z)	coordinate	for	printing.After successful printing of uniform cross-section 3D pats, more complex parts with non-uniform cross-section were printed. These results are shown in Figure 7.

These results are shown in Figure 7. The ball-in-a-cage part shown in Figure 8A demonstrates an example of the ability of CAL to print geometries which cannot be printed by layer-by-layer methods without support material. The octet truss structure in Figure 8C demonstrates the ability of CAL to print complex structures that can only be manufactured additively.948DiscussionThe present work demonstrates a careful design and initial implementation of an additive manufacturing system which can promote a new paradigm in the way such systems are designed and operate. It fosters a shift from the conventional approach of building 3D parts layer-by-layer toward an approach which is designed to address the entire volume of the 3D geometry. This method can present advantages in manufacturing throughput as well as in the ability to print overhanging geometries without the need for mechanical support material.

This method can present advantages in manufacturing throughput as well as in the ability to print overhanging geometries without the need for mechanical support material. To further fuel this shift, there are some modifications to the prototyped system that will be implemented in future work. First, some improvements can be made by the design of a system where the resin volume is static and the optics rotate around it. The advantage of this system design is that the optics can be rotated at a much higher speed than the volume. This arises from the constraint of fluid motion in the resin volume at higher rotation speeds. The only constraints on the speed with which the optics can be rotated are motor rotation speed limits and the minimum frame rate of the optical system. Neither limit has been nearly approached in the current system. Finally, a future implementation of this system could be designed such that projections from all angles are generated simultaneously.

Neither limit has been nearly approached in the current system. Finally, a future implementation of this system could be designed such that projections from all angles are generated simultaneously. Such a system could then achieve true single step printing of complex 3D geometries.AcknowledgementsThis work was performed under the auspices of the U.S. Department of Energy of Lawrence Livermore National Laboratory under Contract DE-AC52-07NA27344. This work was performed under LDRD funding 14-SI-004, and 17-ERD-116. This document has been released under IM review # LLNL-PROC-733997. We would like to thank both the Center for Engineered Materials and Manufacturing at Lawrence Livermore National Laboratory and the Design for Nanomanufacturing group at the University of California, Berkeley. We would also like to acknowledge useful discussions with Robert Panas, Allison Browar, James Oakdale, and Ryan Hensleigh at LLNL as well as with Prof. Laura Waller, Jingzhao Zhang and Prof.

mputer Aided Contour Profiling of High Strength DepositsSriram Praneeth Isanaka1, Amar Bala Sridhar2, Frank Liou1, Joseph W. Newkirk31Department of Mechanical & Aerospace Engineering, Missouri University of Science and Technology, 400 W. 13th St., Rolla, MO 65409, USA2Department of Manufacturing Engineering, Missouri University of Science and Technology, 400 W. 13th St., Rolla, MO 65409, USA3Department of Materials Science & Engineering, Missouri University of Science and Technology, 223 McNutt Hall, Rolla, MO 65409- 0330, USAKeywords: Metal deposition, deposit contouring, coordinate measurement, 3D scanning, curve fitting, contact angle, ripple effect.ABSTRACTAdditive manufacturing processes suffer from the effect of ripples, edge rounding and surface variations. To reduce their effect, ideal process parameters for the laser deposition process were investigated. Also, a new method was identified to analyze deposits by accurately plotting their contours.

To reduce their effect, ideal process parameters for the laser deposition process were investigated. Also, a new method was identified to analyze deposits by accurately plotting their contours. This was achieved through point cloud data of the deposits generated using coordinate measurement and 3D scanning. Curve fitting was performed on the data in Matlab to generate the contours of the deposit. The intercept values, heights, and contact angle of the curves give an indication of the uniformity of deposits and aid in reducing defects.INTRODUCTIONThe Laser aided material deposition at Missouri university of Science and Technology consists of two major subsystems: Laser deposition system (Rofin Sinar 025) and CNC mill (Fadal VMC-3016L). The subsystems share common x-y translator and rotary-tilting table; but have independent z-axis movements and the entire system is capable of 5-axis motion. The laser deposition system includes a laser and a powder delivery system.

The laser deposition system includes a laser and a powder delivery system. The principle of laser deposition is as shown in figure 1. The Rofin Sinar laser with energy up to 1KW is used to fuse powder on to the underlying surface (substrate). The laser delivers heat while the powder delivery system feeds powder at a uniform rate into the melt pool created by the laser. The 5-axis CNC machining makes it possible to directly produce fully-functional parts with a quality surface finish. Its many advantages include: Eliminating the need for support material, crafting complex contours easily, reducing processing time and eliminating profile constraints which are typical with 3-axis milling. Figure 2 shows the Laser Aided Manufacturing Process (LAMP) system during deposition. The Hardness values obtained by the LAMP system are comparable to other manufacturing processes.

Figure 2 shows the Laser Aided Manufacturing Process (LAMP) system during deposition. The Hardness values obtained by the LAMP system are comparable to other manufacturing processes. The major applications of the process include coatings, functionally graded materials, and repair and maintenance of designs prone to crack failure.Figure 1: Principle of LAMPFigure 2: 5 – axis laser deposition systemPAST WORKGraphic Representation of results is essential to understand the numerical behavior of a mathematical model solving a physical problem. The representation substitutes analysis of a mess of large data by a simple visual inspection. Contour plotting of data is one step towards such representation. Contour plots show the same value of a physical parameter in a geometric domain. It is an indispensable method for graphical representation of results.

Contour plots show the same value of a physical parameter in a geometric domain. It is an indispensable method for graphical representation of results. Contour line plots offer visual information that is not obvious from looking at voluminous numerical data.For contour profiling data obtained from CMM or 3d scanning was used. The use of 3d scanning ranges across many fields. It has been used in geometric morphometric analysis of craniofacial shape [1], laser profilometer [2], Feature extraction of concrete tunnel liner [3] and structural modeling [4], documenting data from 3d body scanning [5], art [6] and sculptured surface digitization [7]. It is useful because the laser scanning data simultaneously provides geometric and radiometric information of objects.

It is useful because the laser scanning data simultaneously provides geometric and radiometric information of objects. The geometric data, the x, y and z coordinates of the objects, are recorded by the principle of time-of-flight, which calculates the distance between the laser scanner and the target with the traveling time of the laser or via the phase difference of the laser rays. Through accurate geometric information provided by laser data and high sampling density, precise information regarding the three-dimensional (3D) features of structures can be provided.CMM has remarkably improved the ratio of cost and performance. It has achieved wide popularity as a flexible solution to a myriad of measuring needs. CMM captures the coordinates of points on the work piece surface. It finds its usage more in the post production or quality stage [8]. Our use of the CMM and 3 D scanning is more in the design & development stage.

It finds its usage more in the post production or quality stage [8]. Our use of the CMM and 3 D scanning is more in the design & development stage. We have used measurement tool in the initial stages of our design. This is used to give greater insight into the process involved and will enable in taking decisions for either rapid prototyping or manufacturing.a) Ballingb) Clearc) Sinteringd) High sinteringg) Uneven widthf) Thin track Figure 3: Classification of LAMP powder deposits by visual inspectione) High Aspect ratioThe major process parameters that affect the finish of the deposits made by LAMP are the powder used, powder feed rate, table velocity and laser power. Every powder including stainless steel & titanium to alloys such as tribaloys and tungsten carbide matrices possess diverse physical properties. Therefore the ideal deposition values for the process parameters vary with differing materials.

Therefore the ideal deposition values for the process parameters vary with differing materials. In the past, ideal deposition parameters were identified by visual inspection of the deposits [9]. The classification of powder deposits by visual means in shown in figure 3. High resolution digital images of the deposits are taken and analyzed visually to classify them into clear, sintered, uneven width or balled tracks etc. The use of this method to identify ideal parameters is tedious and time consuming. It is also not very accurate as it involves the viewer’s intuition and is mostly based on past experience.To address these difficulties we identified the need for a more quantifiable method. The major priority of this new method was to improve accuracy during deposit analysis without human error.

The major priority of this new method was to improve accuracy during deposit analysis without human error. The comprehensive data obtained from this method is to be used in process planning software being developed at Missouri S & T for the LAMP process.PROCEDUREAfter identifying the need for a novel method, significant brainstorming led to the proposal of creating 3-dimensional surface maps of our deposits. This was accomplished by generating point cloud data of the outer surfaces of the deposit and the substrate. The data was then post processed using matlab to provide meaningful results. Error identification and verification then followed and applied to the point cloud data to obtain very accurate 3-D surface plots of the deposits. These plots led to the identification of the variations in deposit heights, intercept lengths and edge rounding phenomenon with greater accuracy. A schematic of the procedure is shown in figure 4.

These plots led to the identification of the variations in deposit heights, intercept lengths and edge rounding phenomenon with greater accuracy. A schematic of the procedure is shown in figure 4. The procedure can be broadly classified into 4 stages: data acquisition, post processing of data, accuracy verification, and output.Stage 1: Data AcquisitionStage 2: Post processingCMM or 3‐D scanningPoint Cloud dataMatlab erroneous point removalData slicing, curve fittingFinal Output: 3‐D surface data including variation in heights, intercept widths and edge rounding.Error identification and correctionProof of conceptStage 4: Output Stage 3: VerificationFigure 4: Schematic for obtaining 3-D surface dataStage 1: Data AcquisitionIn an effort to obtain accurate 3-D surface plots, point cloud data of the surface was gathered by using Co-ordinate Measurement Machine (CMM) and 3-D scanning.

Illustrations of the point cloud data obtained by the two methods are shown in figures 5 and 6. The CMM used for obtaining the above data was developed in-house, for use in this project. It has a resolution of 0.0025 inches and generates approximately 2-3 million points per scan. Very little data losses are found while using this process and it provides a complete scan result. The disadvantage with this process is that it is time consuming. It takes approximately 48 hours to obtain 2 – 3 million points (single run). It provides point cloud data in .vtk formats that can be imported into matlab for post processing.a)b)Figure 5 (a & b): Visualizations of CMM point cloud data in MAYA.The other method used for data acquisition is 3-D scanning. For the experiments a NEXTENGINE 3-D desktop scanner was employed. It has a much higher resolution than that of the CMM. This method produces approximately 10 million points per scan which is both an advantage and a disadvantage.

It has a much higher resolution than that of the CMM. This method produces approximately 10 million points per scan which is both an advantage and a disadvantage. Although it ensures larger and better data set, the large numbers of points pose considerable problems while post processing. The advantage of this process is the minimal time required for data acquisition. A complete scan of 10 million points can be made in less than 90 minutes making this process extremely fast and efficient. The major problem with this method is loss of data owing to line of sight issues as exhibited by most 3-D scanners. The scanner acquires only data that is unobstructed in its view. This can be clearly seen in figure 6 (b).

The scanner acquires only data that is unobstructed in its view. This can be clearly seen in figure 6 (b). This method provides data in .xyz format which can also be imported into matlab for post processing.Data lossesa)b)Figure 6 (a & b): Visualizations of 3D scan data obtained using NextEngine scanner.Both the methods are ideal for our process as they generate the required data. Alternatingbetween the two based on the need for speed and accuracy was best option that was proposed. The point cloud data obtained by this method was then imported and post processed in matlab.Stage 2: Post processinga) Point cloud data plotted in matlabb) Point cloud data after junk data pointsremovedd) Point data from single slicec) Data split into individual slicese) Single slice data with curve fitf) Rough contour profile of the deposit.Figure 7 (a, b, c, d, e, f): Procedure for post processing of data (All axes represent dimensional values in x, y & z directions.

All dimensions are in inches)During post processing, the point cloud data with 3 millions points is broken down into manageable sizes of around 50,000 points. Such a broken down file is shown in figure 7 (a). It can be seen that the point cloud contains points which show the surface of the substrate. All points that do not represent the actual deposit in question are considered junk and are filtered out of the file using matlab programming. Such a file that includes only the deposit point cloud is shown in figure 7 (b).During the course of the analysis it was noticed that removing these junk data bring down the number of points by about 50 – 60 %. The data at this juncture contained approximately 20,000 points. Although the clutter had been significantly reduced from the data the number of points, were still too great to draw meaningful conclusions. Therefore we came up with the idea of slicing the data to further reduce the number of points.

Therefore we came up with the idea of slicing the data to further reduce the number of points. The slices were picked at regular intervals and can be used to predict the variations of height and the trend of the data without the need for using the entire data set. Figure 7 (c) shows a data set with a few slices. These slices were then fit with curves to represent the outer shape as shown in figure 7 (d) and 7 (e). A large number of these slices together can accurately represent the outer surface of the entire deposit as shown in figure 7 (f).Figure 8: Output after post processing the data. (Axes represent dimensional values in x, y & z directions. All dimensions are in inches)The greater the number of slices employed the more accurately they will portray the varying trend in the data. Employing slice data helps identify heights and intercepts accurately while also ensuring a small data set to work with.

Employing slice data helps identify heights and intercepts accurately while also ensuring a small data set to work with. The output obtained after the post processing is shown in figure 8.Stage 3: VerificationThe point cloud data obtained by this method contains numerical values up to the eighth decimal point. Although the data seems comprehensive and can predict trends very well, thevalidity of the data obtained was yet to be verified. The next task was to ascertain that the data obtained matched the cross-sectional profiles of the deposits made.During our initial experiments forty eight deposits of tribaloy- 400 were made on 4340 steel by varying the powder feed rate, table velocity and laser power. The deposits are shown in figure 9. To aid in the verification, the deposits were coated with J B weld and the cut to photograph the cross-sections of the deposits. The cut cross-section would now represent a single slice of data obtained from matlab.

The cut cross-section would now represent a single slice of data obtained from matlab. The photograph of the cross-section and the slice data image from matlab were then superimposed to check the validity of the data. Such an image is shown in figure 10.Figure 9: Experimental deposits of tribaloy - 400Figure 10: Superimposed images for verificationIt can be seen from figure 10 that data obtained from matlab and the photographs are a near perfect match. This proved that data obtained was valid and any error in the data was beyond the 4th decimal point which was to be negated by fitting curves to the data. Based on the above results a number of intuitively chosen curves were fitted to the data set. The curves primarily investigated were polynomial curves of multiple orders and moving average curves from 2 – 25 points as shown in figures 11 and 12.

The curves primarily investigated were polynomial curves of multiple orders and moving average curves from 2 – 25 points as shown in figures 11 and 12. Owing to the simple shape of the deposit cross-sections there was no need to investigate more complex shapes such as splines or Bezier curves.Figure 11: Moving average with 6 pointsFigure 12: Polynomial curve of order 2(Axes represent dimensional values in x & z directions. All dimensions are in inches)Cross-sections of deposits can be made in both x and y direction. After our analysis it was found that polynomial surfaces of lower order fit the cross-sections made along the x-axis well enough to depict the outer surface of the deposits.

After our analysis it was found that polynomial surfaces of lower order fit the cross-sections made along the x-axis well enough to depict the outer surface of the deposits. For the cross-sections of deposits made along the y – axis moving average curves with around 25 points and polynomial curves of higher order were noticed to best depict the actual deposit as shown in figures 13 and 14.Figure 13: Moving average with 25 points along y-axisFigure 14: Polynomial curve of order 5 along the y - axis(Axes represent dimensional values in y & z directions. All dimensions are in inches)RESULTS AND DISCUSSIONFrom the final point data and curves, intercept widths, varying height values and edge rounding effects were obtained. Figure 15 show the intercept points of multiple slices with connecting lines which indicate the width.Figure 15: Intercept points and widths as seen from the top of the deposit. (Axes represent dimensional values in x & y directions.

(Axes represent dimensional values in x & y directions. All dimensions are in inches)It can be seen from the figure that the intercept widths vary with deposition. All the 48 deposits made indicated variations to a certain extent. The trend is that at the start and end of the deposition the intercept width is smallest. During the course of the deposition the variations persists, but to a lesser degree than at the start and stop. This a common problem in all powder additive manufacturing processes where controlling the powder particle’s position to accuracy greater than 0.001 inches is still difficult. The method proposed can identify these minute variations and help with process planning. With the above data all problems related to geometric difficulties and post process material removal can be significantly reduced.The method also very accurately plots the height profile of the deposits. Using this, the extent of edge rounding effect due the varying laser heat was analyzed.

Using this, the extent of edge rounding effect due the varying laser heat was analyzed. The edge effect is predominant during start and stops of the deposition and was initially assumed to be equal in size to the width of the laser beam. The new measurement method though indicates that it is 1.5 times the width of the laser beam. Also the ripple effect or waviness in the deposit due the variations in the process parameters was also gathered. Variations in height due to sintering and powder flow velocity have been studied in depth with a much more quantifiable method to generate optimum parameter values. The height variances during deposition are shown in figure 16.StopDuring depositionStartFigure 16: Varying height profile of the deposit with a moving average trend. (Axes represent dimensional values in y & z directions.

(Axes represent dimensional values in y & z directions. All dimensions are in inches)Upon analyzing all forty eight of the tribaloy – 400 deposits using contour profiling we have found that tribaloy – 400 is best deposited using the values shown in table 1. Very fine deposits of tribaloy – 400 were made using the values in the table. An example is shown in figure 17.Table 1: Ideal parameter values for tribaloy – 400 depositions obtained using computer aided contour profilingS. No 1 2 3Parameter Laser Power Table Velocity Powder feed rateValue 800 W 14 inches/min 6 inches/minFigure 17: Sample deposit made using ideal process parametersCONCLUSIONSThe need for a new measurement and analysis method was indentified and computer aided contour profiling proposed and implemented. Point cloud data from both CMM and 3-D scanning was obtained and post processed to get the required results.

Point cloud data from both CMM and 3-D scanning was obtained and post processed to get the required results. Heights profiles, edge and ripple effects and varying intercepts widths have been analyzed for tribaloy – 400 deposits. This has led to the prediction of ideal process parameter values and a change in outlook over the edge effect occurring in the LAMP process. When this data is used in combination with a process planning software it can help inPredicting ideal material and particle sizes. • Predicting geometric complications. • Calculating tolerances. • Reducing post process milling. • Predicting ideal LAMP machine parameter values.This major advantage of this method is its ability to be applied in any application. Rather than using 3-D scanning and CMM as art or quality control tools, we can now use them to obtain very accurate data and analyze. The equipment cost involved in this process is also minimal.

Rather than using 3-D scanning and CMM as art or quality control tools, we can now use them to obtain very accurate data and analyze. The equipment cost involved in this process is also minimal. The flexibility and accuracy of this method can be employed to good effect in anatomical studies, biological sample measurements, conceptual designs of next generation aircrafts and cars, etc. or other places where there is a need to profile complex contours accurately.ACKNOWLEDGEMENTSThis research was supported by the National Science Foundation grants DMI-9871185, IIP- 0637796, and IIP-0822739. Grant from NUTC is also greatly appreciated. Support from Product Innovation and Engineering, LLC, Missouri S&T Intelligent Systems Center, and the Missouri S&T Manufacturing Engineering Program, is also greatly appreciated.REFERENCES1. Hennesy RJ, Kinsella A, Waddington JL.

mputer Aspects of Solid Freeform Fabrication:Geometry, Process Control, and DesignRichard H. Crawford Department of Mechanical Engineering The University of Texas at AustinAbstractSolid Freefonn Fabrication (SFF) is a class of manufacturing technologies aimed at theproduction of mechanical components without part-specific tooling or process planning. Originally used for creating modelsfor visualization, many industrial users of SFF technologies are realizing the greater potentialofSFF as legitimate manufacturing processes for producing patterns and, in some cases, functional.parts. Thus, SFF is becoming an important aspect of the product realization process in these industries.Solid Freefonn Fabrication arose from the dream of "push-button" prototyping, in whichsolid reproductions of three-dimensional geometric models are created automatically under computer control.

Perhaps more than any other class of manufacturing technologies, computer software development has been an integral part of the emergence of SFF. As SFF technologies evolve toward the ability to create functional parts, computer issues gain more importance.This paper discusses three aspects of software design for SFF: processing of geometricdata, global and local control of SFF processes, and computer-based analysis and design for SFF manufacturing. The discussion of geometric processing issues focuses on accuracy and completeness of input models, and the algorithms required to process such models. The interplay between the physics of SFF processing and the desired output geometry is discussed in terms of the development of model-based control algorithms for SFF. These two areas, geometric processing and control, are necessary for the practical implementation of any SFF technology.

These two areas, geometric processing and control, are necessary for the practical implementation of any SFF technology. However, for SFF to realize its potential as an alternative for manufacturing functional parts, engineers must be provided with analysis and design tools for predicting mechanical properties, ensuring dimensional accuracy, choosing appropriate materials, selecting process parameter values, etc. For each of these three different but related areas of software design, the state-of-the art is assessed, contemporary research is summarized, and future needs are outlined.IntroductionSolid Freeform Fabrication (SFF) is a group of emerging technologies for fabricating physical objects directly from computer-based descriptions (such as solid models) of the geometry of the parts. All SFF technologies are enabled by computer hardware and software, from the input geometric descriptions to control of the fabrication machinery.

All SFF technologies are enabled by computer hardware and software, from the input geometric descriptions to control of the fabrication machinery. Clearly, the success of SFF depends on the existence of sophisticated and cost-effective computing equipment and software. As the applications of SFF move beyond visualization models and design verification prototypes into fabrication of functional components, SFF will change from rapid prototyping techniques into legitimate manufacturing technologies. One key to this progress is improvement of the software systems that enable advances in the capabilities of SFF. This paper describes potential opportunities for improvement in SFF software in the areas of geometry processing, process modeling and control, and design tools for SFF.102Geometry ProcessingMost Solid Freefonn Fabrication processes produce parts on a layer-by-Iayer basis. Theprocess proceeds by fIrSt slicing the geometric description of the part into layers.

Theprocess proceeds by fIrSt slicing the geometric description of the part into layers. The slicing operation generates the contours of the part for each layer. The contours are then processed in a manner dependent upon the particular SFF technology. For instance, for Selective Laser Sintering (SLS), the contours are discretized into "toggle points" at which the laser beam must be modulated to produce the desired solid.Faceted Geometry. The form of the geometric description of mechanical parts to beproduced by SLS significantly affects the accuracy of the final part. The current state-of-the-art for most SFF technologies consists of tessellating the surfaces of the geometric model into a mesh of non-overlapping triangular facets. The resulting geometry is transmitted in a standard file formal.. the so-called "STL" file fonnat, established by 3D Systems, Inc.

The resulting geometry is transmitted in a standard file formal.. the so-called "STL" file fonnat, established by 3D Systems, Inc. [1] This fonnat has been adopted by many CAD vendors, is readily available, and is considered adequate for most visualization applications. However, for producing accurate patterns and functional parts, the adequacy of the STL format is unclear. There is a trade-off between the accuracy and the size of the geometric description. Highly non-linear surfaces, such as those that comprise turbine blades, manifolds, etc., must be tessellated into a large number of small facets, resulting in very large data files, and the accuracy of such descriptions is still suspect.In many cases, the tessellation operation itself introduces errors in the model. Tessellation of surfaces with large curvature can result in errors at the intersections between such surfaces (see Figure 1), leaving gaps or "holes" along edges of the part model [4].

Tessellation of surfaces with large curvature can result in errors at the intersections between such surfaces (see Figure 1), leaving gaps or "holes" along edges of the part model [4]. Tessellation of fine features is susceptible to round-off error, which leads to non-manifold models of parts, where more than two facets are adjacent to a single edge, or facets with opposing outer normals meet at a single vertex (see Figure 2). These problems are difficult for slicing algorithms to handle and cause fabrication problems for SFF processes, which require valid solids as input.Algorithms to overcome these problems with faceted geometric descriptions depend on developing richer data structures that explicitly represent the facet and edge adjacencies that are implicit in the STL format This information enables the development of efficient slicing algorithms that take advantage of the explicit adjacency information when constructing slice contours [5, 11, 12].

Rock and Wozny [10] have also proposed an alternative to the STL file fonnat that captures the adjacency information in the exchange file. B~hn and Wozny [4] have developed a method to repair faceted geometric descriptions with problems such as that depicted in Figure 1. Theirtechnique uses the adjacency information to locate gaps between facets. They then use a "shell-elosure" algorithm to add missing facets to the close gaps.Higher Order Geometric Descriptions. The problems with the STL geometryexchange format arise because tessellation is a first-order approximation of more complex geometric entities. An obvious solution to these problems is to exchange higher order geometric entities, preferably the source geometry with which a part is designed. This approach has been adopted at The University of Texas, where an interface based on Constructive Solid Geometry (CSG) has been developed to provide input to SFF processes [7].

This approach has been adopted at The University of Texas, where an interface based on Constructive Solid Geometry (CSG) has been developed to provide input to SFF processes [7]. With CSG the geometry of a part is modeled as a binary tree whose leaf nodes are scaled and oriented instances of primitive shapes (e.g., spheres, cylinders, parallelepipeds, etc.) and whosenon-tenninal nodes are the regularized Boolean set operations union (u), intersection (fl), and difference (-). Evaluating the CSGtree with a geonletric modeler results in an explicit boundary representation of the part's surfaces.103Surface 1Surface 2Figure 1. A gap in the seam at the intersection of two surfaces [4].Common edgeCommon vertexFigure 2. Other examples of non-manifold topology in a faceted SFF model.104Nonnal evaluation of a CSG tree involves computing the intersection between pairsprimitive or intennediate surfaces and performing set classification in three dimensions.

The process is based on evaluation of sliced primitives rather than 3-D primitives. In this process the primitives in the CSG tree are sliced individually, generating a slice for each primitive. The part contour the slice plane is then produced by combining the primitive slices based on the Boolean operations at the non-tenninal nodes (see Figure 3). The contour of the part in a given slice plane is a collection of piecewise continuous curves. These curves are then scan-converted to produce toggle points.net result of this work is a boundary evaluator developed specifically for applications. The approach provides a more accurate, compact part description. Also, all surface surface intersections are perfonned in the 2D slice plane, resulting in considerably simpler and more robust algorithms. For the common quadric surfaces (spheres, cones, cylinders, and ellipsoids), the surface-surface intersection calculations are exact (subject to round-off errors).

For the common quadric surfaces (spheres, cones, cylinders, and ellipsoids), the surface-surface intersection calculations are exact (subject to round-off errors). For higher order surfaces, such as the torus, the method requires an approximation of the slice contour of the primitive. Such an approximation is still more accurate and efficient that the linear approximations obtained from a faceted model. In fact, the method provides a rational basis for approximating the geometry, since it is based on error measures in the slice plane. Details of the algorithm are included in [7].Processing higher order geometry for SFF processes offers several advantages overexchange of faceted geometric descriptions. Generally speaking, files containing higher order infonnation will be smaller and more accurate than comparable faceted geometry files. Also, many of the problems that result in non-manifold geometric information in faceted descriptions can be avoided.

Also, many of the problems that result in non-manifold geometric information in faceted descriptions can be avoided. For example, because the CSG slicer described above perfonns the geometric modeling operations, no explicit boundary information is needed in the input file. Instead, it is incumbent upon the SFF geometry processor to ensure that the results are realizable for the particular technology that is used tofabricate the part. Potential problems in the slicing operation can be solved because more infonnation is available about the intended geometry of the part; thus, higher order descriptions are easier to troubleshoot when necessary. Finally, when approximations are necessary for the given input geometry, the approximation process is driven by the particular SFF technology rather than by generic criteria meant to satisfy the requirements of many SFF technologies.

This provides a rational basis for approximating the geometry when necessary.There are disadvantages to higher order geometric data exchange as well. First, there is nosingle geometry fonn that is satisfactory for all applications. There are many different geometric descriptions that are used in product design, each with different requirements for a slicing algorithm. Designers of commercial SFF processing software will have to make compromise decisions about which geometric fonns to support or risk losing potential customers from geometric coverage.• Also, because the geometric inputismorecomplex, algorithms for processing the geometry are more complex as well. Finally, interchange standards must be developed for higher order geometric descriptions. While these standards are under development, they are still subject to change until agreed upon.Sources of Layered Geometry. Many potential applications layer-based formats.

These are just three examples of the potential applicatioIls forareas other than product engineering.each of these cases the data was presented to.............'-'u.J.u""' as layers.105sphere1 : center at (0, 0, 0) and unit radius sphere2 : center at (1 , 0, 0) and unit radius sphere3 : center at (0, 1, 0) and unit radiusA slicing ~ sphere3 >SP~1 ~ere2circle2CSG TreeSliced Treeyycircle1circle2circle2Before applying BooleansAfter applying BooleansFigure 3. CSG slicing and set operations for SFF [7].The geometric processing was not particularly sophisticated in any of these cases. Forinstance, in each of the cases cited above, the thickness of the data slices did not coincide with the thickness required for the SLS process (typically 0.005 to 0.01 in). To compensate for this, the SLS control software either skipped data layers (for thinner layers) or replicated layers (for thicker layers).

To compensate for this, the SLS control software either skipped data layers (for thinner layers) or replicated layers (for thicker layers). To achieve higher accuracy, however, more sophisticated strategies must be developed and adopted by the SFF software. The laser digitizer software described in [13] actually interpolates between slices to realize the required resolution for the SLS process. Similar schemes must be developed for processing cr and confocal microscope imaging data.Both the cr data [8] and the microscope data [2] consist essentially of raster images of the physical objects being imaged. In each case, the data provides a measure of the relative density of the material at each pixel in the imaging plane. However, SFF processes such as SLS require boundary information rather than interior density data.

However, SFF processes such as SLS require boundary information rather than interior density data. For these studies, then, the raster images were frrst converted to contour images of the sample slice in each plane by specifying a minimum threshold density below which the data are ignored. For these studies, the threshold value was selected manually and varied from layer to layer. However, to realize the full potential of1C6tbuild IdirectionConstantthicknessVariable slice thickness4. Effect of variable slice thickness on part accuracy.models with SFF, threshold values must be detennined automatically. Algorithmsuv','I.IJI.'l.AUUUAjO, optimal threshold values are needed for preprocessing layer-based SFF inputuse of layered data can expand, data interchange standards must be established. are no standards to specify the fonn of layer-based geometry for exchange amongne(~eSl;ltv for accommodating such data in SFF processing.

are no standards to specify the fonn of layer-based geometry for exchange amongne(~eSl;ltv for accommodating such data in SFF processing. In the examples cited n,'"-.A.''' presented the data to the SLS machine in the fonn of contours [5]. The other two........ '.---<""above, present their data as raster bit maps. These examples suggest that"".n."'AAI,...'•• ji", ... standards should be developed: one for contour data and one for raster data. In any U,",d.'I-LAA" for processing the input geometry, i.e., interpolation or thresholding as described left to the particular SFF technology to ensure that maximum part quality can beProcess Modeling and ControlFabrication has the potential for producing accurate, structurally sound"UJ!.A,lAvA,I,;)Jl\,.i'JlAu.JL solid versions of objects. To develop accurate patterns and functional parts with however, requires in-depth basic research to understand the correlation between~ ..... ~.~ .. ~.'"AJlAv,.... AAL4.. AJl....

To develop accurate patterns and functional parts with however, requires in-depth basic research to understand the correlation between~ ..... ~.~ .. ~.'"AJlAv,.... AAL4.. AJl.... u.JL properties and geometric accuracy of the final part with respect to the SFF processThese relationships must be captured in computer models that can be used toprocessing.enable real time control, these models must represent a compromisemodeling the physics of the SFF process and intelligent use of geometricinfonnation to approximate optimal processing patterns.While process planning for SFF is considerably reduced compared to conventionalfabrication technologies, there are several considerations which will require reasoning about the geometry of the final part. Scaling and orientation of the part within the work space of the SFF machine have a significant impact on the efficiency of the process.

Scaling and orientation of the part within the work space of the SFF machine have a significant impact on the efficiency of the process. Aside from other factors, the part should oriented in a manner which minimizes the number of layers. Other factors, however, may override this consideration. For instance, tolerances tend to be directionally U'-'&.l'-'UU'-'AA". Likewise, the mechanical properties of the fmal part willdepend upon its orientationThese issues require geometric reasoning on both global and local scales.VU.'VAAJlAAjO, will indicate the best part orientation within the workspace of the SFF machine. to considerations of the geometry of each layer to determine scanning and"-IIJ.,"-1UUJIreasoning .... lltt,Q'M... C' that maximize geometric accuracy and mechanical properties of the part.107Scanning patterns. Many SFF processes build each layer of a part by raster scanningthe powder bed along a single axis, as indicated in Figure 5.

Many SFF processes build each layer of a part by raster scanningthe powder bed along a single axis, as indicated in Figure 5. For geometries such as this, a unidirectional scanning pattern will result in a large number of very short scanning vectors. A mutidirectional scanning pattern, such as that indicated on the right in Figure 5,·will result in a smaller number of longer scanning vectors. Longer vectors reduce the errors associated with laser toggling.·transients andrepositioningofthe.laserbearn,resultitlginbigberpartaccuracy.The figure implies thatpreferred scanning patterns aretberesult ofgeometrya!one;however, other factors rnay outweigb geometry. ForinStal1ce'rnechanica!properties witbina layer may be dependent upon the scanning direction; beat transfer considerations and the creation of thermal gradientsrnay indicate preferred scanning patterns; contour scanning of the layer boundaries may reduce· aliasing artifacts. and result in better·surfacequality.

and result in better·surfacequality. These considerations must be incorporated into algorithms for determining arbitrary scanning patterns for SFF processes.Contour Scanning..• Recentresearch atUT Austin has focused on developing an optimallaser control system forscanninga!ongcurves[15]. The scanning algorithm and associated hardware maintain constant laser power density by simultaneously controlling laser speed and laser power. The goal is to produce Parts with better surface resolution than can be obtainedfrom vector scanning the part contour (see Figure 6). This research emphasizes the need for higher level geometric descriptions for SFF process. The algorithm uses information about the the curvature of the contour to determine appropriate laser parameters to achieve the desired power density. A polygonal approximation of the contour, such as that obtained from slicing a faceted part model, is not accurate enough to support this control scheme.Process Modeling.

A polygonal approximation of the contour, such as that obtained from slicing a faceted part model, is not accurate enough to support this control scheme.Process Modeling. Advances in control of SFF processes will depend on developing a better understanding of the physics of each process. For example, the process control program for SLS administers three areas: laser control, control of powder delivery and leveling, and control of the environment within the machine. Input to the module consists of a suitable geometric description of the part layers (at this time, scan lines and laser toggle points) and settings for process parameters, including material properties, bed temperature, gas flow rate, scanning speed, laser power, beam diameter, and scan spacing. The current SLS process control software is built under the assumption that these parameters will be constant for an entire part.

The current SLS process control software is built under the assumption that these parameters will be constant for an entire part. However, initial investigations indicate that fabrication of metal parts will require local control of laser beam parameters, allowing these parameters to change from layer to layer or even within different areas in a given layer. Such physically-based scanning is depicted conceptually in Figure 7, where a part layer has been divided into several regions based on part quality predictions from a physical model of the process. A scanning pattern is then generated for each of the simple regions, again based on predictions from a process model. Such a scanning scheme could be precomputed off-line before the part is fabricated, or it may updated on a layer-by-Iayer basis, with input about the status of previous layers from sensors. The key to solving this problem is development of physical models of the process.

The key to solving this problem is development of physical models of the process. While initial models of the SLS process have been developed [9, 14], they are not structured to support real-time process control of this kind. The solution to this problem will likely require a combination of hardware and software development108Raster ScanningDirectional ScanningFigure 5. Directional scanning versus raster scanning.Vector ScanningContour ScanningFigure 6. Contour scanning.109Figure 7. Physically-based scanning.Design for SFFSuccessful design of functional parts produced by SFF will require consideration of thematerial and mechanical properties in conjunction with their geometry. These properties are directionally dependent due to the layer-by-layer fabrication process. For certain SFF processes, such as SLS, the mechanical behavior of parts is further complicated by the phase change inherent in the process.

For certain SFF processes, such as SLS, the mechanical behavior of parts is further complicated by the phase change inherent in the process. Therefore, standard techniques for modeling structural behavior of mechanical parts will have to be modified and enhanced to account for these factors. Development such computer-aided design tools will allow designers to create parts that are optimized processes.Physical models of SFF processes provide a starting point for developing SFF tools. However, such models are focused on determining the consequences process parameters on the properties of the final part. For design, the inverse solution is ne(~o~a. For desired properties of the final part, the designer must know the appropriate process parameters. Answering such questions will require reformulation of the modelingSFF processes offer the promise of providing manufacturing capabilities that are not realizable by other techniques.

Answering such questions will require reformulation of the modelingSFF processes offer the promise of providing manufacturing capabilities that are not realizable by other techniques. One such possibility is selective material property distribution within the part With conventional material removal processes, the bulk mechanical properties part are determined by the stock material chosen, aside from any surface treatment that is "V&JAA\,.-Y a post-process. With SFF technologies the potential exists for the mechanical continuously within the part. Again, to realize this possibility, design tools are designer in determining optimal material distribution.

With SFF technologies the potential exists for the mechanical continuously within the part. Again, to realize this possibility, design tools are designer in determining optimal material distribution. Project MAXWELL, a joint University of Michigan and Carnegie-Mellon University, is developing mathematical ......""........."4 ...,"'.:# for concurrent design of shape and material composition for mechanical componentsto110tosuch as this will become increasingly important as SFF techniques are employed for manufacturing functional parts rather than prototypes.ConclusionOne key enabler of Solid Freefonn Fabrication technologies is computer software in geometric modeling and process control, without which most SFF processes would be difficult to implement. However, many opportunities exist for improving the performance of SFF processes by improving the software, thereby broadening the application of SFF from producing models and prototypes to functional part manufacturing.

This paper provides an outline of some of the issues in software development in the areas of geometric modeling, process modeling and control, and design for SFF. Progress in anyone of these areas, however, is dependent upon improvements in the other two areas. These interdependencies are depicted in Figure 8. To realize the potential of SFF, research must continue simultaneously in all of these areas.Geometry • Quality depends on process parameters. • Determ ined by intended design function.process Control • Orientation, laser parameters depend on geometry. • Design function determines desired properties, which depend on processing.D!i!sign • Limited by available processes and capabilities. • Auxiliary design requirements (supports) determined by geometry.Figure 8.

MPUTER GENERATION OF METAL COMPONENTS BY SIMULTANEOUSDEPOSITION OF MOULD, CORES AND PARTMasoud M. Mohebi, ShoufengYang and Julian R.G. Evans Department of Materials, Queen Mary, University of London, Mile End Road, London, E1 4NS, UKAbstractA new solid freeforming method based on co-delivery of mould powder materials and part powder materials using vibration-controlled, dry powder valves is presented in this paper. Thin layers of stainless steel powder are delivered to the forming area according to the cross-section of the CAD file to produce the component. Mould powder which has low sinterability is delivered to the non-forming areas of the same layer. All powders are delivered by computer-controlled, acoustic powder valves. The flow rate and switching of the valves provides the composition and shape control during fabrication. The stacked layers of loose powder are then sintered in a conventional furnace. The mould materials are removed after sintering.

The stacked layers of loose powder are then sintered in a conventional furnace. The mould materials are removed after sintering. This method avoids the high thermal stress problem in selective laser sintering, avoids high capitalisation, makes use of conventional furnaces and allows for the incorporation of three dimensional function gradients. Test pieces including step wedge and Spierpinski’s cube were fabricated. Advantages, limitations and problems are discussed.IntroductionSolid freeforming (SFF), reviewed by Pham and Gault [1] among others, is defined as the construction of a three dimensional shape by point, line or planar addition of material without confining surfaces other than a base [2]. In existing SLS, the same material is used in the forming areas and non-forming areas. The material in the non-forming areas is used to support the next layer. Powder in the forming areas is fused by laser scanning.

The material in the non-forming areas is used to support the next layer. Powder in the forming areas is fused by laser scanning. Normally the materials in the non-forming areas are difficult to recycle due to the thermal history. The selective laser sintering process involves wide temperature ranges, rapid cooling and oxidation problems. Furthermore the generation of residual stresses is an inevitable consequence of one layer solidifying upon its predecessor [3, 4]. We therefore devised a method of assembling a complex shape capable of having 3D compositional gradients and for which conventional sintering could be applied. As with SLS, there would still be restrictions on compositional range and on compatibility of powders. These are set by the demands of co-sintering, but the procedure would be much simpler. It would allow for compaction if required.

These are set by the demands of co-sintering, but the procedure would be much simpler. It would allow for compaction if required. Furthermore, advanced fast sintering methods such as spark plasma sintering (SPS) [5] or induction sintering [6 ] would be able to co-sinter a wide range of compositions and may improve compatibility.The emergence of the second generation of solid freeform fabrication methods requires techniques of dispensing multiple materials on a powder bed before laser sintering in SLS or binder printing in 3DP. Various powder mixing and dispensing methods have been studied in the490past few years, such as discrete multi-powder positioning [7], removal of unsintered powder by vacuum suction [8], powder dispensing robot [9], powder ejected by piston and weighting by resonance measurement [10], vibration method [11,12, 13, 14, 15, 16], and electrostatic method [17, 18, 19].

Rock and Gilman [20, 21] describe a system in which two types of powder, one acting as the part powder and one as a support powder that also defines the shape of the part are deposited in each layer. The method of achieving this is not recorded and the description refers throughout to a ‘deposition effector’ which is not defined. Pegna et al. [22, 23], inspired by the ancient sand-painting art, presents a method of making complex shapes from construction materials that cannot be realised by casting. A similar approach has been used to produce conventionally sintered components, which can include 3D functional gradients, by using acoustic powder feeding of the part and mould powders concurrently [24]. In each layer, the forming area is made of the part powder or mixtures of powder of computer-defined composition and the non-forming areas consist of a mould powder which can later be removed.

In each layer, the forming area is made of the part powder or mixtures of powder of computer-defined composition and the non-forming areas consist of a mould powder which can later be removed. In this way, it is not necessary to sinter or melt each layer as in SLS, but rather, the whole assembly can be sintered in a conventional furnace such that the mould powder retains the shape of the part. Thus strategies for fully three dimensional functionally graded components based on powder technology are beginning to emerge. In this paper, the fabrication of test pieces including step wedge and Spierpinski’s cube (Menger’s sponge) were demonstrated.Experimental DetailsThe powders were gas atomized grade 420 stainless steel (16-63μm, ex Osprey Metals, Neath, Wales), alumina (BACO Chemicals, Batch: C731 UNG Al2O3, BA Chemicals Ltd, Buckinghamshire, UK.), copper (63-212μm, ex Osprey Metals, Neath, Wales) and graphite (GS150-E, GrafTech, UCAR Carbon Co., USA).

The properties of powders were listed in Table 1.Table 1. Properties of powders*PowderTruedensity/ kg.m-3Tap density/ kg.m-3Loosedensity/ kg.m-3Angle of repose/ (cid:101)Hallflow meter test /s (50g)Stainless steel Graphite Alumina *Errors are 95%CL80002100-2200 39704600(cid:114)200530(cid:114)100 1480(cid:114)3004120(cid:114)200500(cid:114)100 1350(cid:114)30029(cid:114)349(cid:114)3 33(cid:114)229.5(cid:114)2.5No flow 88.3(cid:114)4.8The powder delivery device consisted of glass capillary tubes of inside diameter 300–660 μm. This diameter range of capillary tubes gave on/off switching and flow rate control for the test powders used. Acoustic vibration was provided by 25mm miniature bass drivers. The cone centres of these were connected to the ends of the tubes by 1mm diameter and 40-50mm length rods.

Acoustic vibration was provided by 25mm miniature bass drivers. The cone centres of these were connected to the ends of the tubes by 1mm diameter and 40-50mm length rods. The vibration was supplied from a 2×30 watt amplifier by sound generated using software and491converted into analogue signal by a DA card on a personal computer. The detail of the powder dispensing device is shown elsewhere [25]. A 3D table was driven through 6K4 controller and motion planner software interfaced through labview. The flow rate of powder was measured using a four place balance interfaced to the PC. The reading of the balance was recorded every second. Rectangular graphite crucibles (89mm(cid:117)33mm(cid:117)33.5 mm, inside dimension) were machined from stock (GrafTech, UCAR ltd, Sheffield, UK) and provided with a close-fitting base capable of sliding in the ‘Z’ direction to provide the building platform.

The schematic diagram of the equipment and the powder delivery device are shown in figures 1 and 2, respectively.Mould powder nozzlePart powder nozzlePowder bedCrucibleMould powderPart powderCore powderZ AxisPistonX-Y TableFigure 1. Schematic diagram of powder delivery into a crucible fitted with building platform.All the experiments were performed at 20-25°C and under 30-60% RH, a region known to provide little variation in flow properties (Matei et al. 1973). The samples were partly sintered in flowing Ar+4%H2 at 1000-1100(cid:113)C and infiltrated with bronze at 1100(cid:113)C. Sintered parts of stainless steel were cut, polished and inspected for carborization based on ASTM E407-99 for etching for detection of carbides. Electrolytic etching using a solution of ~10wt% ammonium persulphate in water at 6 V for 30 s was carried out (specimen as anode).492CapillariesSpeakersGraphiteCrucibleFigure 2 The equipment for dual powder delivery.

Six capillaries were installed of which 2 were used in this experiment.Results and discussion1. Compatibility of mould and part powders.In selective laser sintering, an entire layer is placed on the platform and the part powder is selectively laser fused while powder in the non-forming areas remains in the loose state. However in the present method, both the part powder and the mould powder remain in the loose state before entering a conventional sintering furnace. The part powder is either sintered or partly sintered at a temperature where the mould powder remains loose. The compatibility of the mould powders and part powders needs to be considered.

The part powder is either sintered or partly sintered at a temperature where the mould powder remains loose. The compatibility of the mould powders and part powders needs to be considered. The requirements are a) the mould powder must have a higher sintering temperature than the part powder; b) the two powders should have low reactivity; c) the mould powder should be capable of easy and total removal after sintering; d) the powder should be sufficiently free-flowing to allow its flow to be arrested and regulated by the acoustic valve. An additional requirement of core powder compare with the mould powder is that it should be able to “collapse” to avoid hot cracks analogous to the requirement of the cores in sand casting. From the experimental results, the mould materials fulfilled the extra requirement as core materials so we adopted same materials as mould and core materials in this paper.

From the experimental results, the mould materials fulfilled the extra requirement as core materials so we adopted same materials as mould and core materials in this paper. The operation of the valve and powder characteristics that allow successful metering and dispensing have been discussed elsewhere [14,26].493Both coarse alumina powder and graphite powder were tested as the mould/core material. Loose copper powder and alumina powder were loaded into an alumina crucible with a thin paper partition which was carefully removed before sintering. These powders were also loaded into a compression die with a thin paper partition which was carefully removed and the assembly pressed at 80 MPa. The interface of non-pressed copper is fairly smooth while in the pressed one, the compaction behaviour, which can be assessed by instrumented die pressing of each powder [27] , is not a perfect match and as a result, the interface between the two powders is slightly distorted after pressing.

The non-pressed and pressed copper/alumina samples were then sintered in a furnace at 1050(cid:113)C for 3 hours with a reducing atmosphere (Ar - 4%H2). The alumina powder was not sintered at all. The copper could be easily cleaned although there were some alumina particles embedded in the surface. The sintered samples were examined using SEM and backscattered electron mapping of aluminium (Figure 3). The pressed sample was sintered to quite high density while the non-pressed was just weakly bonded. In the pressed copper there was a thin layer of entrapped alumina power. The problem presented by the alumina powder entrapped at the surface is that it is difficult to remove the last few grains from sintered parts. Some of the alumina particles appear to be mechanically locked at the surface.

Some of the alumina particles appear to be mechanically locked at the surface. Since alumina has a hardness of 2050 kg/mm2 (Knoop indenter), attempts to finish by conventional tooling are impeded although shot blasting with an abrasive powder may be effective.abNotPressedPressedFigure 3: (a) SEM and (b) backscattered electron mapping of aluminium atoms of 2 sintered specimens of pressed (left) and non-pressed (right) cupper powder. Bright line in (b) indicates the entrapment of alumina powders in the pressed copper part. Contamination of alumina in body section of the samples is due to spread of alumina powder throughout the porous specimen in the wet cutting processes for SEM.Graphite was also tested a s a mould material. Loose stainless steel powder and graphite er were loaded in an alumina crucible with a thin paper partition which was carefully powd removed before sintering.

Loose stainless steel powder and graphite er were loaded in an alumina crucible with a thin paper partition which was carefully powd removed before sintering. Then the sample was sintered in a reducing atmosphere (Ar - 4%H2) at 1100 (cid:113)C for 1-2 hours. The low loose density and flowability of graphite gives problems when more layers are deposited; the graphite powder has greater compaction ability than steel powder. After multi-layer deposition, the denser steel powder on top of the graphite tends to compress it494more than the neighbouring steel. This may cause distortion of previous layers as the steel powder sinks into previous graphite layers. The main disadvantage of graphite is the possibility of carbon pick-up by ferrous alloys, which is shown in Figure 4. The depth of the carburization is 100-200μm from the interface.

The main disadvantage of graphite is the possibility of carbon pick-up by ferrous alloys, which is shown in Figure 4. The depth of the carburization is 100-200μm from the interface. No carburization was detected in the central region of the samples.abFigure 4: Optical micrograph of partially sintered stainless steel with graphite powder as mould after electrolytic etching to reveal the carbides. The bulk samples do not show carburization (a) but it is detected on the surface adjacent to graphite powder (b).here are several approaches for removing the mould powder. It can be achieved simply by T anical detachment, by a gas phase reaction that does not affect the part or by dissolution in amech suitable solvent or reactant for which the part material is neither soluble nor reactive.2. Flow rate and composition controlhe flow rate of the powder can be controlled over a wide range by varying the tube diameter, T vibrat ion frequency and vibration amplitude.

Flow rate and composition controlhe flow rate of the powder can be controlled over a wide range by varying the tube diameter, T vibrat ion frequency and vibration amplitude. At low frequencies and amplitudes, domes form in the tube and flow stops. The valve can thus provide flow rate control as well as switching. Previous experiments [26] on the effect of tube diameter on flow arrest indicated that if the ratio of tube radius R, to particle radius r, was such that 2<R/r<5, flow arrest occurred when vibration ceased. A model was established and an equation was derived which provides a reasonable analytical prediction of the experimental flow rate given the complexity of the flow regime [28].

A model was established and an equation was derived which provides a reasonable analytical prediction of the experimental flow rate given the complexity of the flow regime [28]. In the open-capillary valve, flow rate approximates to a function of reciprocal amplitude and the gradients show that flow rate doubles for only 20% increase in tube diameter.enefitting from a wide range of flow rate control by varying the tube diameter, vibration B frequ ency and vibration amplitude, the powder dispensing device can be constructed either for single materials with several tubes of different diameter or for tubes holding different materials. A dry powder dispensing device with 6 valves was constructed and shown in Figure 2. More valves495could be mounted with optimized spacing arrangement. With this multi-valve dry powder deposition device, on-line materials mixing and pattern deposition was achieved [25].

More valves495could be mounted with optimized spacing arrangement. With this multi-valve dry powder deposition device, on-line materials mixing and pattern deposition was achieved [25]. Colour sand paintings with 5 different colour sands have been made to demonstrate the feasibility of multi materials deposition [29].3. Test piece fabrication The materials could be eith er vector scanned or raster deposited [30]. To avoid the errors in the on /off control and improve the dispensing efficiency, the layers were vector scanned to fill the areas (Figure 5). The quality of the build is influenced by the fill strategy. In the early stages of this investigation, the areas were filled with rectangular spirals or zigzags. Examination of the building sites showed that particles were piling up at the turning points as shown in Figure 5 (c) and (d). The diagonal lines seen in these steel powder beds are due to pile-up at the right angle bends.

The diagonal lines seen in these steel powder beds are due to pile-up at the right angle bends. This is caused by the low acceleration and deceleration rate at the turning points, which was needed to avoid shaking the three-axis table and disturbing the powder assembly. While the displacement speed of the table at the turning point is lower than the speed during straight line deposition, the flow rate of the powder is constant. Although it is possible to reduce the flow rate at the tuning point, the response of the flow rate is slow compared with that of the 3D table and the programming of the vibration parameters is complex. This looks acceptable for one or two layers but during multi-layer deposition, this defect can be magnified into seriously uneven surfaces, as can the defect caused by the track distance.

This looks acceptable for one or two layers but during multi-layer deposition, this defect can be magnified into seriously uneven surfaces, as can the defect caused by the track distance. To avoid this defect, curved turning points and uniform XY table speed were used wherever possible [24].o fill a powder blanket properly, the track width and track distance in the same material T zone and the track distance between adjacent mould and part materials determine subsequent distortion of the pattern. Improper filling gives an uneven powder bed surface which accumulates into serious inaccuracies after multiple layers of powder deposition. The track width was affected by powder properties, flow rate, nozzle to powder bed distance, and vibration conditions. A study on track width is described elsewhere [25]. The track distance between mould powder and part powder is extremely important in providing for a smooth and well-defined internal part surface.

A study on track width is described elsewhere [25]. The track distance between mould powder and part powder is extremely important in providing for a smooth and well-defined internal part surface. A large track distance between mould and part materials produces a valley as illustrated in Figure 6(a). Conversely, a small track distance produces a hill as illustrated in Figure 6(b). This kind of defect in the powder bed is difficult to observe in the initial stage of deposition but becomes magnified into a serious problem during the fabrication of a vertical wall. An example of the hill effect is shown in Figure 6(d). At the top surface of the sample, the edge adjacent to the mould powder was piled up due to a small track distance between the part and mould powders.496(a)(b)(c)(d)Figure 5 (a) (b)Vector scanning strategies for area fill and (c) (d) co-filled graphite and steel powder blanket on SiC grinding paper.

In (a) and (b), the dark lines represent the track of steel and the light lines represent the track of mould/core powder. The lines linking two filled area only represent the motion route; the powder valve is closed.In order to find the ideal part-mould powder track distance, a step wedge test piece was designed. The track distance was varied at each step from 0.8mm to 1.1mm. The step test piece was infiltrated with bronze and sliced vertically to examine the steps as shown in Figure 7. From right to left the track distance of the step is 0.8, 0.9, 1.0 and 1.1mm. Notice that the ‘hill’ effect is present on the 0.8mm step and the ‘valley’ effect on the 1.1mm. The 0.9mm step offers the best track distance. A similar situation develops at the perimeter of each layer adjacent to the crucible where an improper distance causes hill or valley effects around the wall of the crucible.

A similar situation develops at the perimeter of each layer adjacent to the crucible where an improper distance causes hill or valley effects around the wall of the crucible. However, because of imperfect fit between crucible and moving base or piston, there is always slight leakage of powder at the crucible/piston gap and this means that the design must be such that part powder does not extend to the crucible wall.497(a)(b)10mm(d)(c)Figure 6 Illustration of defects in the cross section of powder bed. (a) Valley caused by large mould-part powder track distance. (b) Hill caused by small mould-part powder track distance. (c) Ideal powder bed made from optimum mould-part powder track distance. (d) An example of the hill defect (indicated by arrow).(a)(b)Figure 7 Step wedges fabricated by steel-alumina co-deposition. Fired at 1250(cid:113)C and infiltrated by bronze. From right to left the track distance of the step is 0.8, 0.9, 1.0 and 1.1mm.

Fired at 1250(cid:113)C and infiltrated by bronze. From right to left the track distance of the step is 0.8, 0.9, 1.0 and 1.1mm. Hill effect was present on 0.8mm step and valley effect on the 1.1mm. The 0.9mm is the best track distance.A test piece of a part/mould/core powder co-deposition is shown in Figure 8. The test piece is a slab of 27(cid:117) 27(cid:117)9mm3 with a vertical square middle hole of 9(cid:117)9(cid:117)9mm3. Both the part/mould (Figure 8 (a))and part/core interface (Figure 8 (b))has flat surfaces. The presence of the core materials did not introduce hot cracks. Ripples presented in the part/mould and part/core interfaces indicated the layered manufacturing. The smooth wall of the graphite crucible gave a smooth but carburized part/crucible surface. A non-ideal part/crucible wall track distance also gives a hill498defect at the edge shown in Figure 8 (a).

A non-ideal part/crucible wall track distance also gives a hill498defect at the edge shown in Figure 8 (a). To avoid the carburization near the wall of the crucible, mould powder was used to insulate the part/crucible thereafter.10mm5mmPart/cruciblePart/mouldPart/core powderinterfacepowder interfaceinterface(a)(b)Figure 8 Test piece for a part/mould/core powder co-deposition showing (a) the vertical wall achieved on the part/mould powder interface and (b) the vertical wall of the part/core powder interface.A test piece of the first order of Menger’s sponge was fabricated (Figure 9). The Menger’s sponge was selected as a test piece to show the capability of complex core fabrication. The horizontal holes were slightly distorted due to mould powder-filling defects (mould powder has higher overrun than the part powder due to the high flowability of the alumina).

The horizontal holes were slightly distorted due to mould powder-filling defects (mould powder has higher overrun than the part powder due to the high flowability of the alumina). The defect could be eliminated in further investigations of the filling track parameters of the mould powder.10mm(a)(b)Figure 9. (a) 3D model of Menger’s sponge and (b)the test piece fabricated.ConclusionsA new solid freeforming method based on co-delivery of mould powder materials and part powder materials was demonstrated in this paper. Both coarse alumina powder and graphite powder were tested as the mould/core material. The interface between the dissimilar powders were499investigated. The problem presented by the alumina powder is the entrapment at the surface while the disadvantage of graphite powder is the carburization and powder compaction. The flow rate and switching of the valves provides the composition and shape control during fabrication.

The flow rate and switching of the valves provides the composition and shape control during fabrication. Optimization of track width and track distance as well as filling strategy eliminated the fabrication defects. Test pieces including step wedge and Menger’s sponge were fabricated with acceptable smooth interfaces. Further investigations are needed on the optimization of remaining fabrication parameters.AcknowledgmentsThe authors are grateful to the Engineering and Physical Sciences Research Council forsupporting this work under Grant No GR/S74164.References1 D.T. Pham and R.S. Gault , A comparison of rapid prototyping technologies, International Journal of Machine Toolsand Manufacture, Vol. 38, 1998, 1257-1287.2 B.Y. Tay, J.R.G.Evans, and M.J. Edirisinghe, Solid Freeform Fabrication of Ceramics, International MatererialsReview, Vol.48, 2003, 341-370.3 A. Vasinonta and J.

Evans. On the rate of descent of powder in a vibrating tube. Philosophical Magazine, 85 (10),2005, 1089-1109.29 S. Yang and J.R.G. Evans. Automated sand painting. To be published.30 J.L. Johnson, Principles of computer automated fabrication. Palatino Press, 1994, p11.501.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperComputer integration for geometry generation for product optimization with Additive ManufacturingT. Reiher*, S. Vogelsang* and R. Koch*Direct Manufacturing Research Center (DMRC) and Chair of Computer Application and Integration in Design and Planning (C.I.K.), The University of Paderborn, Mersinweg 3, 33098 Paderborn, GermanyAbstractDesigning parts for additive manufacturing (AM) offers a broad range of geometrical and functional potentials. On the one hand the manufacturing technology offers the possibility of manufacturing highly complex freeform shapes, often referred to as bionic shapes. By use of these, perfect force fluxes without stress risings due to imperfect notches are realizable, getting the most value of used material.

By use of these, perfect force fluxes without stress risings due to imperfect notches are realizable, getting the most value of used material. On the other hand these complex structures require a reliable geometry representation in compatible CAD-files. Conventional CAD systems were developed to generate geometries that are manufacturable with conventional machining. These are not capable of representing the high complex designs for AM. Especially for geometries generated by CAE like from topology optimization the conventional CAD systems fail to take advantage of the combination of CAE and AM. This paper explains why there is a lack of compatibility of well-known CAD systems with the potentials of AM. Therefore the AM-side of the problem is described by showing some potentials of AM and the need of high complex structures for this manufacturing technology.

Therefore the AM-side of the problem is described by showing some potentials of AM and the need of high complex structures for this manufacturing technology. For the other side of the problem conventional methodologies for geometry representation of CAD systems are described and their limitations with regard to AM are worked out. Finally a voxel based geometry representation is presented as a solution for computer aided geometry generation of high complex AM–structures.IntroductionNowadays additive manufacturing (AM) has evolved to a reasonable direct manufacturing technology for actual components, made of a variety of materials. Parts are created by material addition layer by layer combining two dimensional layers (2D) made from small volume elements (1D) to three-dimensional shapes (3D). Therefore the manufacturing is conducted tool less, as the material is deposited, molten and solidified in place without a shape-giving tool like a cutter or a mold.

Therefore the manufacturing is conducted tool less, as the material is deposited, molten and solidified in place without a shape-giving tool like a cutter or a mold. Thereby AM can be differentiated from conventional manufacturing technologies by two core characteristics: the “1D–2D–3D-principle” and tool less manufacturing, taking into account that “1D” represents small volume elements – voxels – what is very important for the later described geometry representation methodology. [RK16; RK15] The layer wise and tool less manufacturing enables a broad range of geometrical and functional potentials between which the potential “complexity for free” temporary had become something like a mantra for selling AM as a promising technology [HHD05]. Beside this a lot of other potentials as monolithic design and function integration or flexible semi-automatic production of slightly different designs in one manufacturing shot are named.

Beside this a lot of other potentials as monolithic design and function integration or flexible semi-automatic production of slightly different designs in one manufacturing shot are named. [GRS10; RK16] No matter whether these potentials are way more easy to realize with AM instead of conventional manufacturing many of them have been discussed years ago [EKH+07]. Crucial for a successful903manufacturing of parts and especially for exploiting the potentials of AM is a well-designed digital representation of the latter part. Not merely for “complexity for free” but rather most of the AM- potentials and particularly for computer generated lightweight components, very complex geometries arise. Conventional CAD-tools are often doomed to fail for these complex geometries as their basis was developed many years ago for replacing 2D-drawing and the clear aim of generating geometries possible to be made with conventional machining.

Thus there is a need for new digital geometry representation methodologies in CAD tools, especially developed for high complex freeform designs as a vital link of computer aided geometry generation and additive manufacturing. Therefore in the following an explanation and comparison of requirements coming from AM and capabilities of CAD tools will be given to work out a new voxel-based principle to enhance the overall process.Existing CAD representation schemesIn over 50 years development of CAD-systems many different methodologies for geometry representation were invented. All of them are following the same aims of describing and capturing geometrical objects and their attributes with a strong focus on geometry visualization. Therefore suitable data structures have to be used and for a long time complex geometries had to be approximated by more simple ones [Sha01].

Therefore suitable data structures have to be used and for a long time complex geometries had to be approximated by more simple ones [Sha01]. Following [Sha01] there are three general rules mandatory for each engineering language in terms of solid modeling to guarantee and maintain their integrity:“Any constructed representation should be valid in the sense that it should correspond to some real physical object;Any constructed representation should represent unambiguously the corresponding physical object;The representation should support (at least in principle) any and all geometric queries that may be asked of the corresponding physical object.”Thus, there is something like a common sense in feasibility, usefulness and correctness postulated. Validity, correctness and correspondence to a physical object is one of the major challenges in terms of high complex geometry representation and the direct evolution from digital to physical product subsistence.

The most direct conversion from digital to physical is boon and bane in the same time, as there are barely fallback levels to overcome failures in the digital geometric data. When thinking about efficient generation of geometric data for AM this has to be kept in mind especially in contrast to design for conventional machining like milling where little freeform surfaces but mostly flat or concentric surfaces for turning are cut. In a mathematical understanding there is the difference of not-, simple- or double curved elements. For conventional manufacturing the double curved elements are seldom seen as this is way more complex to produce than not-, or simple curved elements, while for AM the additional effort is circumstantial. [Ada15] Methodologies for solid modeling can be classified via the representation scheme and the modeling processes. Thereby the representation scheme is the technology that is actually representing the geometry and enables the user to see a geometry on screen.

Thereby the representation scheme is the technology that is actually representing the geometry and enables the user to see a geometry on screen. In comparison the modeling processes describe the way the geometry is generated by means of defining the location and relationship of each point and edge that will be used in the representation. In the beginning of computer aided design as digital support for engineers, only simple 2D- drawings as direct substitution for hand drawings were available. Based on these very simple digitalization of manual work, over the years many different approaches for representing geometric objects have been invented. The great number of different schemes can be categorized in three categories as shown in Figure 1: direct, indirect and parametric representation schemes.

The great number of different schemes can be categorized in three categories as shown in Figure 1: direct, indirect and parametric representation schemes. [Fra09]904Figure 1: Representation Schemes in solid modelingThe category of direct representation schemes is specifically marked by representing a solid volume by direct representation with volumetric elements. Thus single mistakes like missing small elements do not destroy the overall geometry. For indirect schemes like a polygon mesh the absence of one single element ruins the closed surface so that no entire volume is surrounded and the solid cannot be detected. Furthermore, with regard to the geometry generation for additive manufacturing, direct representation schemes based on norm cells or voxel grids enable simple local editing like smoothing when single elements are added or removed.

[RK15] On the contrary for other schemes like CSG (Constructive Solid Geometry) or norm cells with low resolution it is difficult to represent complex freeform shaped surfaces. Voxel based schemes like spatial occupancy enumeration are used for example in computer tomography scans (CT) to save the scanned information for each volume element. Those scans enable the representation of very high complex geometries but require very big storage systems.905One of the first invented representation schemes in CAD systems is based on simple primitives used with Boolean operations for addition and subtraction to create more complex geometries. Figure 2 gives an example for such a complex geometry generated by the combination of very simple volumes and one cutting plane. This technology is called Constructive Solid Geometry (CSG) and until today is one of the most underlying and used principles. Studies show that a lot of parts are representable only with these features.

Studies show that a lot of parts are representable only with these features. All of these are easily manufacturable conventionally with respect to undercuts or the like. A survey from the evolving time of CAD has shown that “63% of all the parts could be handled with a CSG system based on only orthogonal block and cylinder primitives. A larger class of primitives provided a natural description of over 90% of the parts. This indicated that CSG is therefore a good fit for a CAD system in that sort of environment because most mechanical parts seemed to be relatively simple.” [SRE76; Ago05] Another representation scheme is the Cell Decomposition Representation. This is what is mostly used for finite element modeling as the geometry is decomposed into plenty small elements like triangles in two-dimensional cases or tetrahedra for three-dimensional models [Ago05].

Thus the cell decomposition is a versatile principle as the use of triangles is possible as well for indirect representation schemes (B-REP) if a volume is represented only via a closed surface. In this case the tetrahedra are small volumetric elements whose combination results in a direct volumetric geometry representation. Near to the cell decomposition the Spatial Occupancy Enumeration represents a geometry with small uniformly sized cells. In this case the volume is represented by small squares in 2D (pixel) or small volume elements in 3D (voxels). These voxels are ordered in a fixed grid and all of them that are inside the volume are marked as volume elements whose coordinates are explicit. As already written bevor, the main problem of this representation is the big amount of data that has to be stored. Thereby this scheme was not used much for CAD systems. Though today the computing power and memory have increased that much that this scheme gets more and more important.

Thereby this scheme was not used much for CAD systems. Though today the computing power and memory have increased that much that this scheme gets more and more important. [Ago05] Due to its capabilities and the voxel-based approach that is near to a special interpretation of the AM technology this scheme is very usable in the context of geometry generation for AM [RK15; RLJ+17]. Thereby later on this scheme will be discussed again in detail.Figure 2: Constructive Solid Geometry [Sha01]The next category is the indirect representation schemes. The main characteristic feature in this case is, that the geometry is not represented by any volume but more by the edges or the surfaces one can see from the outside and the volume is not explicitly represented. In the beginning of the development of three-dimensional CAD systems the Wireframe Representations have been the most usable way of representing a geometry because it is easy and fast to compute and display.

Thereby the vertices of a geometry are stored in a database and then connected with edges. These edges can be displayed and thereby a wireframe of the model is shown. Unfortunately for these representations it is hard to decide which edges are in the front or behind the rest of the model and one cannot get cutting view. On the other hand for complex geometries it is very fast to display and one can see through the part to identify inner structures. Today most CAD systems still enable such a view. [Ago05]906In the group of surface based indirect schemes the main important one is the Boundary Representation (B-REP) scheme. In this scheme the geometry is represented only via its surfaces. If the surface of a geometry is fully closed, the inner volume can be interpreted as a solid. The overall surface of the geometry can be interpreted as a relationship between objects and certain graphs.

The overall surface of the geometry can be interpreted as a relationship between objects and certain graphs. Hereby the surface consists of n faces which are based on n edges that are always connecting 2 vertices. Inside the graph the vertices can be used by arbitrary edges that can be used for arbitrary, but reasonably two surfaces. [Ago05] B-REP as well as CSG is one of the basic representation schemes that have been used from the early beginnings of CAD system development. B-REP thereby is capable of being combined with many other schemes. Figure 1 names the Polygon and the Sweeping Representation as explicit schemes, but can be interpreted as or at least easily translated in a B-REP as well. Sweeping is a scheme where a set of lines is “swept” along another line and thereby encloses a volume. Thus this representation corresponds naturally to the way many mechanical parts are manufactured like sheet metal systems, turning or milling [Ago05].

Thus this representation corresponds naturally to the way many mechanical parts are manufactured like sheet metal systems, turning or milling [Ago05]. Polygon meshes again are well-known from FEM analysis as mentioned for the Cell Decomposition already. Thereby the Surface of a three-dimensional geometry is divided in small finite elements, mainly triangles or squares. The difference is that the two-dimensional surface can be of any high complex freeform shape but has to enclose a volume so that the inner area can be interpreted as a solid. The decomposition of the boundary representation with triangles to enclose a solid volume is as well what is known from the .stl data format that is used as a de facto standard for additive manufacturing. Here one of the main problems of indirect representations can appear.

Here one of the main problems of indirect representations can appear. If one of the great number of triangles, easily up to a million for big complex structures, is flipped around to indicate the wrong side as inner volume or is missing at all, the inner volume is not fully enclosed and cannot be interpreted as solid. This fault cannot appear for direct representation schemes. The last category of representation schemes is the parametric category. Mainly the parameterized primitive instancing, the parametric and feature-based modeling and the implicit representation can be named here. [Fra09] As the category name already is indicating, the main characteristic is the strong influence of parameters. Because of the inferior importance for the topic discussed in this paper these are only explained in short. Parameterized Primitive Instancing is the representation scheme used for part libraries as for screws.

Parameterized Primitive Instancing is the representation scheme used for part libraries as for screws. Thereby certain primitives are fully described with parameterized measurements that are often depended from each other. All single objects are generic primitives which are members of a family of objects, distinguishable from the others by a few parameters. Thus the user can easily change one of the parameters and thereby call another family member. The main drawback of this technique is that not all new geometries can be designed as the family already defines the possible variants. Furthermore the definition of these families is very complex and thus only feasible if a broad usage, as for bolts, is expected. [Ago05] Parametric and Feature Based modeling is what we mostly see in CAD systems and is based on the use of parametric shapes associated with attributes like length, width, radii etc.

[Ago05] Parametric and Feature Based modeling is what we mostly see in CAD systems and is based on the use of parametric shapes associated with attributes like length, width, radii etc. This modeling is often used in combination with CSG representations but with the difference, that there is a clear focus on the parametrization of the used primitives. Thereby a linkage of the model with databases as for variant management is possible.Figure 3: A boundary representation Graph [Ago05]907The last representation scheme is the implicit representation. Hereby the geometry is defined implicitly to consist of all points that satisfy a specified condition. This condition can be a mathematical function and each point describable from this function is a part of the geometry. Thereby complex geometries can be designed with only a small mathematical description.

Thereby complex geometries can be designed with only a small mathematical description. [PAS+95]Existing modeling techniquesAs we now have seen several possibilities to store digital geometries we need to discuss modeling techniques to design these geometries. Some of the modeling techniques can be used for different representations as well as some representations can be filled with different modeling techniques, while for some there are clear relationships. Figure 4 gives an overview of different modeling techniques, clustered in polygon based, curve based and digital sculpting techniques. In general the choice of the right modeling techniques is dependent on the intention what to do and the used representation scheme. Thereby the polygon based modeling techniques are nothing for manual geometry generation. As a complex geometry consists of a plethora of polygons these need to be generated automatically.

Thereby the polygon based modeling techniques are nothing for manual geometry generation. As a complex geometry consists of a plethora of polygons these need to be generated automatically. A definition of each of the vertices and the conducting combination to polygons would be very time consuming and only seems to be feasible for very easy geometries. Thus the polygon based geometries always are generated on basis of other geometry definitions. One possibility therefore is on basis of a point cloud. This cloud can be the result of a 3D-Scan of an existing geometry like from 3D-cameras or tactile measurement arms. Also a definition via mathematical defined points with curves and an appropriate division could be feasible. Another possibility for modeling a polygon based geometry is on a surface mesh on basis of an already existing surface made with any other modeling technique.

Another possibility for modeling a polygon based geometry is on a surface mesh on basis of an already existing surface made with any other modeling technique. A common approach for smoothing geometries gained from any voxel based scanning methodology like CT-scanning is the Marching Cube Algorithm which divides the voxel grid and approximates the stair-stepped geometries to much more smooth geometries on polygon basis.Figure 4: Modeling Techniques overview908Polygon meshes are widely spread in digital geometry representation. The big advantages of a very simple geometry definition and simple data formats in combination with very fast rendering, saving and loading have brought polygons to a standard in graphics processing and 3D-printing. Furthermore most FEA tools make use of polygons for surface meshing. Apart from that of course there are some drawbacks like as already mentioned an active design by defining the coordinates of each vertex is nearly impossible.

Apart from that of course there are some drawbacks like as already mentioned an active design by defining the coordinates of each vertex is nearly impossible. In addition it is impossible to represent curved surfaces really accurate. There is always a triangulation error so that a circle never is perfect round but consists of many straight lines with little angle difference. Thus a high resolution for smooth surfaces lead to a high amount of data. At least it is possible to create meaningless geometries as if a surface is interrupted no solid can be interpreted. In case of 3D printing there have been quite good tools developed to identify those mistakes and correct them automatically. Nevertheless one have to keep in mind that in the end each geometry seen on screen is rendered on basis of polygons and then converted in to 2D-pixels! The second modeling technique branch are the curve based modelings.

There is quite a different to the polygon based models though both techniques can be used for the surface based representation schemes like B-Rep, Polygon and Sweeping. The main difference is that any curved things via polygon based modeling are mapped discontinuously while the curve modeling is capable of representing complex geometries continuously. The simplest one to understand is the direct linkage of modeling technique to representation scheme for the geometrical primitives and CSG, as each of the simple primitives is set up with mathematically clear defined formulas like a circle, cylinder etc. Only the combination generates more complex structures. Though this is hardly capable of representing real freeform elements like bionic looking structures, as these designs are not based on the simple geometric primitives. Nevertheless a careful use of only proper defined primitives like of the so called r-set guarantees a correct model without geometrical mistakes.

Nevertheless a careful use of only proper defined primitives like of the so called r-set guarantees a correct model without geometrical mistakes. [Ago05] Most CAD-Designs start with sketching some 2D representations and then doing some extrusions, sweeping or the like. This is a reasonable extension of 2D design and often is near to conventional manufacturing like milling, turning or extrusion. Over the years these techniques have been implemented very well so that they work well without errors and are easily to combine with other techniques. But if a complex smooth surface shall be designed feasible of taking advantage of the AM- potentials these simple CAD features come to their end. For very smooth surfaces shapes need to be designed by the use of mathematically describable continuously differentiable curves like Splines. Splines in general are two- or three-dimensional lines drawn on a plane or in space whose curvature might change as often as necessary.

Splines in general are two- or three-dimensional lines drawn on a plane or in space whose curvature might change as often as necessary. By combination of many of these mesh can be drawn in space to represent a surface by curves. If we assume only four splines defining each edge of a squared surface this conforms perfectly to the definition of one surface of a Boundary Representation scheme (B-Rep) where only some further surfaces are needed to enclose a volume and form a solid. Later-on basissplines (B-Splines) and non-uniform rational b-spline (NURBS) were developed to define complex shaped surfaces with curvatures in more than one direction. These surfaces are described mathematically precise and thereby capable of representing complex technical geometries.

These surfaces are described mathematically precise and thereby capable of representing complex technical geometries. The generation of these surfaces is easy to manage for geometry systems where the designer can keep track of each surface and the connection to the adjacent surfaces.909Figure 5 shows an example for a part where a manual creation of numerous NURBS for representing a complex topology optimization structure has failed. The combination of over 300 single, handmade NURBS surfaces resulted in a vast number of failures not only can be seen in unconnected surfaces but also in missing G² continuity and thereby smoothness. For G² continuous surfaces the second derivation of the surfaces curvature needs to be continuous.

For G² continuous surfaces the second derivation of the surfaces curvature needs to be continuous. Furthermore NURBS should be capable of enabling perfect smooth junctions of struts and surfaces as fatigue notch factors can be reduced drastically with the right notch geometry this possibility is not easy and nearly impossible for the shown overall geometry complexity. The manual generation of freeform surfaces by use of several splines combined to a surface by something like a mesh is possible. The complex surface transitions in areas of strut-combinations however is manually not possible as a bunch of several surfaces is needed to represent the complexity in those areas. Thereby for curve based modeling with NURBS for setting up a B-Rep model similar to polygons automatic algorithms are needed to support the user. [RK15] Summarizing curve based modeling in general is capable of designing high complex geometries and represent them reliable in the digital world.

[RK15] Summarizing curve based modeling in general is capable of designing high complex geometries and represent them reliable in the digital world. Nevertheless complex combinations with struts are hardly made manually and the process is error-prone. Perfect notch factors as well as G² continuity is hard to realize. Furthermore the designer needs to know how the structure shall look like and it easily can get very complex to adjust single elements after being designed.[DRK15]. Take advantage ofFigure 5: Example for bad geometry retransition by manual NURBS creation [RK15]The third branch are quite new modeling techniques: digital sculpting which is especially capable to overcome the hurdles of freeform design, developing new ideas during design process and making small changes on freeform elements after they have been initially designed.

These techniques are taking advantage of many already mentioned principles but are combining them in a new way to enable to work with the material like in real-life with clay.910The displacement principle, better known as subdivision (Sub-D), sets up surfaces on basis of a cube whose eight corner points as well as the twelve edges are used as control points and the surface in between is the simplest gravitated geometry with a low but equal weighting on each point and edge is a perfect sphere. With maximum weight on each point and edge, the surface is that much pulled to these control elements, that a cube is represented. The underlying principle does not focus on a manual design of new products but more is a method of representing a smooth surface by decompositioning it with polygons step by step as shown in Figure 6. This underlying algorithm was mainly developed by Edwin Catmull and Jim Clark, resulting in the “Catmull-Clark subdivision surface”.

This underlying algorithm was mainly developed by Edwin Catmull and Jim Clark, resulting in the “Catmull-Clark subdivision surface”. The great advantage of subdivision surfaces for modelling are high complex, but faultless, smooth surfaces. Each generated surface is G² continuous what seems to be perfect for both stress equality and fluid dynamics in channels. The systematic refinement of the single decompositions enables the design of big, smooth elements as well as highly complex and very detailed elements in one shot. Furthermore even very complex subdivision surfaces do not require a high amount of storage because only the control points are saved. Even the export in conventional CAD systems is very easy as a conversion into standard NURBS surfaces is no problem. The other main representative for digital sculpting is the voxel based approach. For this approach the geometry is set up by a vast number of small voxels in a clear structured voxel grid.

The other main representative for digital sculpting is the voxel based approach. For this approach the geometry is set up by a vast number of small voxels in a clear structured voxel grid. Thereby it is easily possible to add and remove elements and thereby design a geometry as in real-life with clay. Advanced software tools allow the user even to interact via a haptic interface to really model the part as in real-life with well-known clay-modeling tools. One of these tools is Geomagic Freeform©. Here either an already existing geometry can be transferred into a voxel based model and adapted or the user can start with a block and design a new part. The transformation of a voxel based model in a standard CAD geometry requires some work but is possible. One often used approach is the already mentioned “marching cube algorithm” that works very well on the clear defined regular voxel grid. Additionally quad meshes are easy to generate and retranslate to NURBS.

Additionally quad meshes are easy to generate and retranslate to NURBS. The main disadvantage of both technologies is the lack of a parametric description of geometries. Whilst in general this is technically possible, nowadays it is barely used. This might be a reason why these technologies are mainly found in arts. In engineering areas for the design of technical products it can be less found, except for special usage like texturing of surfaces, what is often made by voxel technology. The handling of these technologies is unfamiliar for most engineers and especially for Sub-D barely comparable with any other method. Furthermore for simple, perhaps easy parametric describable, geometries the needed storage is comparably high but changes for high complex geometries.towardsthese.

Furthermore for simple, perhaps easy parametric describable, geometries the needed storage is comparably high but changes for high complex geometries.towardsthese. TherebyFigure 6: Subdivision of a cube [Rom-ol]What is a complex structure (...good for)?When talking about complex structures realized with AM two questions arise: what characterizes a complex structure and what is the benefit of such structures? Generally a complex911system is defined as a system in which many different parts interact with each other composed from the Latin words “com” (together) and “plex” (woven). Thereby a complex structured geometry consist of many different structure elements which interact with each other. This is not an all-new thing for additive manufacturing, as on the one hand entire products, like the assembled car or plane, have been developed since years. On the other hand single parts of these assemblies have been optimized for its purpose as well.

On the other hand single parts of these assemblies have been optimized for its purpose as well. Thereby one single part can consist of several “features” in a CAD system combining volume elements with holes, radii, ribs and chamfers. AM is now a key enabler to increase this complexity as every single geometry element of a part can be adjusted to interact optimal with the others. Transferred into a geometry this implies that surfaces are not forced to be flat or curved in explicit radii with defined ends but can freely merge very smooth into each other with several overlying curvatures and without discontinuities. These surface combination areas can be optimized for optimal force flux without stress risings due to notch effects and thereby follow renowned bionic principles [MB03; Mat06].

These surface combination areas can be optimized for optimal force flux without stress risings due to notch effects and thereby follow renowned bionic principles [MB03; Mat06]. Especially for geometries that are not handmade but designed with the aid of advanced computer algorithms like topology optimization where a vast number of single elements like big and small struts, stiffening shear plates, interfaces, holes and sponge or bone-like areas are combined to one complex system, a vast number of joining and dividing surfaces have to be designed. That comes to the conclusion that for a pretty good design in terms of lightweight design for optimal material usage, each of these surfaces and combinations need their curvature to be at least described by a third grade function and continuous in the second derivation as long as it is no interface region to adjacent parts.

This implies that two curved surfaces are not only connected by a tangential combination but with the same tangent plane and same normal curvature in a common point resulting in perfect smooth transitions. This is called curvature – or third-level parametric continuity (G²). [Har03; Ago05] With conventional CAD systems such shapes are barely generable by hand in a high number per part. The representation of such surfaces is done by NURBS that need many control points. Automatic functions are needed, as described later on.912MotivationNone of the beforehand described methodologies is capable of representing the desired complexity and functionality of geometries alone. Furthermore the use of computer integration for semi-automatic geometry generation as on basis of topology optimization results require advanced methodologies of geometry generation, representation and modification.

Some of the existing methodologies are made for arts or for gaming and capable of representing complex geometries but fail for the use in actual mechanical components as mentioned before. Most components are only small pieces in bigger assemblies and thereby need some clear defined interfaces for mounting, connection and force transmission. Figure 7 shows a simple lever for such a combination of parametric elements (brown) and freeform optimized geometries around (grey). When designing a new product with freeform surfaces it is necessary to design these elements as parametric defined features like for drill holes or bearing seats. A logic connection needs to be established and possible to be represented and changed in both parts. Conventional CAD methodologies are powerful in modeling such exact defined parametric and easy conventionally manufacturable geometries but have a lack in the representation of high complex freeform surfaces.

A new process, maybe including new data formats for 3D files need to be established.Figure 7: Example for simple lever as a combination of freeform and parametric features and representation via quad meshThus a CAD file containing a geometry taking advantage of the AM potentials has following requirements: - - - - - -automatic geometry generation and optimization shaping of highly complex structures consideration of AM-restrictions compatibility with parametric features fit for the future: multi-material production information content easy variability, adaptability and customizabilityThese requirements will be described and assessed for their usability in the following chapter.

Furthermore already solution approaches will be shown.Methodology for computer integration in geometry generationThe development of a lean method of designing CAD geometries taking full advantage of the potentials of additive manufacturing thereby can be broken down to finding solutions for the six requirements defined above. First point thereby is the automatic geometry generation and optimization. In this case we shorten up the different possibilities of inventing geometries and focus on FE-based technologies like topology and shape optimization and their subsequent steps. Other advanced methodologies913could be a definition based on mathematical functions definition. As these technologies are very well for generating complex arts, it gets hard to define actual mechanical components only with these technologies.

As these technologies are very well for generating complex arts, it gets hard to define actual mechanical components only with these technologies. For topology optimization a finite element (FE) model is set up including all boundary conditions and outside influences like occurring loads, constraints, heat transfer or the like. These are acting upon a virtual volume called design space in where material is allowed but might be reduced if not needed for withstanding the loads while keeping in mind boundaries like maximum displacement, Eigen frequencies or maximum weight. Thus structural parts can be optimized regarding their structural integrity or material consumption. As the standard procedure works on standard finite element models the results are barely directly usable. While for conventional manufacturing the results needed to be redesigned and adapted to the manufacturing technique anyway, no one has been bothered about imperfect optimization results.

While for conventional manufacturing the results needed to be redesigned and adapted to the manufacturing technique anyway, no one has been bothered about imperfect optimization results. Nowadays, AM enables the manufacturing of arbitrary complex shapes so that one can imagine that it is desirable to generate results in a quality sufficient for sending directly into manufacturing. Of course one should never trust the optimization results without sufficient validation and for high value parts it might be reasonable to adapt the results. For simple parts though, a mainly automatic process could reduce environmental impact by material usage reduction while only low effort is acceptable. A feasible approach based on the use of a voxel based software already was shown in 2015 by [RK15]. There the Topology Optimization results are smoothed and adapted after the optimization process to gain smooth and manufacturable designs by converting the results into a voxel representation.

There the Topology Optimization results are smoothed and adapted after the optimization process to gain smooth and manufacturable designs by converting the results into a voxel representation. Afterwards it is transferred back to a CAD representation by generating NURBS patches on the surface, supposedly by use of the marching cube algorithm and quad meshing as already mentioned above. Second requirement is the shaping of high complex structures. On the one hand this aims at the general generation and imagination of complex structures as only these are really economical feasible of being manufactured additively towards other technologies. Additive manufacturing of simple structures mainly is not feasible as traditional techniques are more reliable and less expensive than AM and should be used for geometries they are capable of producing.

One example for AM-specific structures could be the use of lattice structures for inner geometries or making use of the freeform capabilities by creating perfect smooth radii and reduce the notch factors in parts. Both might as well be possible to generate with topology optimization, despite nowadays there is only questionable possibilities of integrating simulation of lattice structures in topology optimization. Thus on the other hand these complex structures do not only have to exist, they even have to be manageable by Simulation, CAD and slicing systems. Again the lattice structures are easy to manufacture and there are high benefits of these structures but is hard to be designed them as the high amount of surfaces to be constructed and displayed on screen is too much for most computer systems. So there is a need of a file system that is capable of replacing these geometries internally in a more manageable thing or better a representation scheme that is capable in general.

So there is a need of a file system that is capable of replacing these geometries internally in a more manageable thing or better a representation scheme that is capable in general. Again a voxels representation might be a solution, as for a given resolution there is nearly no difference on how complex the represented structure is. Nevertheless a very high resolution leads to a big amount of data that might cause other problems that might even be worse. A more detailed discussion on the question of voxel capabilities towards need for storage will be shown in the following. In a third point the geometries have to consider AM-restrictions. AM often is said to be capable of producing any shape easily. In some points this is, especially for metal AM, not correct. As for other manufacturing technologies there are a lot of design rules that have to be considered for a successful print job.

In some points this is, especially for metal AM, not correct. As for other manufacturing technologies there are a lot of design rules that have to be considered for a successful print job. This starts with the need of support structures to channel off the heat from the melt pool to the build platform below and to keep the parts in place during build job. But furthermore the size of a cross section per layer is limited, as the higher the cross section the higher914the internal stresses that might lead to a distortion of the part. [Ada15] These are just two very rough rules of a big bunch of them. So for a worthy CAD file a design check is mandatory to ensure printability. Thus a check of cross section area size, wall thickness and the like has be done before. If this should be integrated in the topology optimization this might come to a big effort of checking the part during each optimization iteration.

If this should be integrated in the topology optimization this might come to a big effort of checking the part during each optimization iteration. As well in this case a voxel grid might enforce the optimization process as it is way more easy to determine cross sections and wall thickness inside a structured voxel grid than on basis of spline-based or triangulated geometries as for those a lot of computation per layer is needed. The voxel grid directly can answer with the number of voxel per layer.Figure 8: Design Methodology based on voxel representation for regaining [RK15]When considering a voxel based geometry the next point of multi-material production information content is very easy to cover, as well. Thinking about AM as a voxel based manufacturing technology theoretically it is possible to print each voxel with a different material, perfectly adapted to the local requirements.

Thinking about AM as a voxel based manufacturing technology theoretically it is possible to print each voxel with a different material, perfectly adapted to the local requirements. At the current state multi-material printing, especially for metals, is only in research and not usable for real part manufacturing. Though, some ceramic processes are already capable of using different ceramics in one process. A voxel based representation scheme easily would be able to include the material properties per volume unit. As there are a lot of potentials and reasons for using a voxel based representation scheme for CAD files taking advantage of all AM potentials, there are still several drawbacks as already mentioned before both from a programming view which will be discussed later on as well as from a logical view.

A real worthy CAD file still needs compatibility with parametric elements to be digitally connectable with other parts as in assemblies and for lean CAD-CAM coupling for the mechanical post processing. As neither polygon models as .stl files nor voxel based models are capable of this, there seems to be a need for a combined representation scheme which will still include standard CAD representations as for example CSG or B-REPs with NURBS. Even more a conversion from voxels to NURBS seems to be mandatory as parts developed and ready for print need to be usable in standard CAD system by standard trained, conventionally very experienced engineers.915Assessment of a voxel based approachThinking about a voxel based approach for making CAD files feasible to produce high value AM parts taking advantage of the full potentials of AM seems to reasonable.

However there are some reasons beside the parametric features, why this kind of geometry representation scheme for engineering has not become a standard by now. For explanation of the hurdles and possible solutions, in the following a detailed view on the voxel based representation scheme with focus on memory usage, representable complexity and modelling techniques will be given. First, as already mentioned above, the memory usage for geometries represented by voxels raises very fast and seems to be not feasible representing a big, complex geometry. If complex freeform geometries shall be represented smoothly by voxels, a very high resolution is needed to overcome stair stepping effects. This is the same as for two dimensional computer graphics represented by pixels. Furthermore using a voxel grid there is a need for having a storage place for each voxel inside the grids space. Thereby the number of voxels easily can raise to very high amounts.

Furthermore using a voxel grid there is a need for having a storage place for each voxel inside the grids space. Thereby the number of voxels easily can raise to very high amounts. A comparison of the two main considerable representation schemes in terms of additive manufacturing and complex designs is given in Figure 9.Figure 9: Comparison of memory consumption for a simple geometry in voxel and polygon representation schemeThere a simple circular geometry is shown represented by pixels and lines as two-dimensional equivalent of voxel and polygons. One can see, that there is a big difference in the actual shape regarding different resolutions with a very bad representation for low resolutions. Nevertheless, as the graphs are well indicating, the voxel representation starts with only small amount of memory for low resolutions but then is raising very fast. On the other hand, for low resolutions the polygons are needing much more memory but then is raising slower than for voxels.

On the other hand, for low resolutions the polygons are needing much more memory but then is raising slower than for voxels. As indicated by the different color of the representations this is, among others, due to the fact that the voxel need to store each single point, while the polygons are only representing the boundary. This equals the categorization of the representation schemes as above where the voxel are called direct representation in comparison to indirect representation via polygons. Though, looking at the memory consumption when keeping into account the complexity of the geometry, there is a big advantage of voxel representation identifiable. Figure 10 shows three different geometries: a simple circle, slightly complex shape and a high complex bone structure represented in voxel, polygon and NURBS. The following comparison of memory consumption for voxel and polygon than shows the big advantage of voxels for high complex geometries.

The following comparison of memory consumption for voxel and polygon than shows the big advantage of voxels for high complex geometries. As each voxel inside the grid needs to be stored,916there is no difference in memory consumption immaterial how complex the geometry represented is while for polygons the memory consumption raises continuously with increasing complexity.Figure 10: Comparison of memory consumption for different complex geometries in voxel, polygon and NURBSThus one can see that indeed the starting memory amount is quite high in the beginning but is equaled for more complex geometries. Furthermore the extremely simple logic behind the voxel grid allows a lot of compression. An impressive study on compression of voxel based geometry representation was done in [VMG16]. By use of very advanced compression algorithms including an octree approach it is possible to store an entire aircraft in only a small amount of memory.

By use of very advanced compression algorithms including an octree approach it is possible to store an entire aircraft in only a small amount of memory. Considering the size of a standard metal AM machine like the SLM 280HL with 280*280*360mm³ and a layer height and thereby resolution of 50µm, a great number of voxels are needed to represent the overall build chamber on layer resolution. Though, these are still less than from the examples in [VMG16] where need less than 100 MB of storage. Thus it seems feasible to store a full build chamber on a resolution higher than printable in a reasonable file size. If this is really working during geometry generation and adaptation has to be proved by further work. The next point is that voxel based representations furthermore are easily capable of enabling geometry manipulation even in high resolutions. For example generating an offset of a geometry is shown in a comparison towards a surface based representation like .stl in Figure 11.

For example generating an offset of a geometry is shown in a comparison towards a surface based representation like .stl in Figure 11. One can easily see, that at the bottom of the groove an intersection of polygons or self-intersection of the surface is expectable that needs to be filtered out by additional algorithms to generate a valid geometry. In comparison to that for a voxel representation the offset generation algorithm is developed very easy and a self-intersection is not existent, as already existing voxel just are existing and there will invalid geometry be generated that needs to be cleaned up. The same benefit can be transferred to other geometry adaption algorithm as Boolean or slicing operations as well.917Figure 11: Comparison of generating an offset with polygon and voxel based representationGenerating offset geometries is a very strong method for the retransition of topology optimization results.

An example is shown in Figure 12 where different steps are compared to retranslate a very simple topology result into a smooth geometry. It can easily be seen that first creating a positive offset that is than smoothed and in a third step taken back with a negative offset generates a very good surface. In this case, there has no manual work has been done. After the rough optimization taking three minutes, the voxel offset and smoothing tooks only one minute. Thereby one can see how fast this voxel based method is.

After the rough optimization taking three minutes, the voxel offset and smoothing tooks only one minute. Thereby one can see how fast this voxel based method is. Furthermore for very high complex parts like the upright, manual effort except in detailed areas would not be feasible and thus an automatic approach is needed.Figure 12: Using offset geometries to generate valid smooth geometry retransitions of topology optimization results918The last task is to guarantee an easy and fast conversion from existing NURBS data out of existing CAD assemblies into a voxel based generation and for the conversion back again to keep standard systems still compatible. Therefore Figure 13 shows how this process can be done nowadays with a polygon based interim stage. As can be seen, the continuous edge from beginning is triangulated by polygons in red. As well it has to mentioned that the beforehand existing additional information regarding the circular bore hole is no longer existing.

As well it has to mentioned that the beforehand existing additional information regarding the circular bore hole is no longer existing. On basis of the polygons it is easily to set up voxels, or in this two dimensional figure, the pixels. After assuming any changes in the geometry there is the back transition by use of a marching cube algorithm triangulating the pixel / voxel back to a polygon based representation. Using further algorithms for auto surfacing, these polygons can be transferred back to the NURBS again. Unfortunately in this process the information on the circular bore hole is still lost. Thus a good surfacing algorithm needs to store the information of the bore hole and include this information when retransferring.Figure 13: Conversions from NURBS via Polygon to Voxel and backFor a real efficient use the interim step of polygons should be avoided and the NURBS need to be transferred into voxel directly.

Theoretically this should be possible but still needs to be proved and developed. First steps on developing a software including a voxel based topology optimization and transitions out of CAD files and back again to enable the voxel based representation methodology to get a usable tool have already been made at Paderborn University. Further work on still open questions and for compressing the voxel representation in feasible sizes still needs to be done.919Conclusion and outlookThis paper has given a detailed description of existing geometry representation and modeling schemes for the generation of high complex CAD designs and the requirements coming from additive manufacturing on these geometries. Based on this a methodology for the generation of high complex CAD designs feasible for additive manufacturing was broken down to six requirements which than has been assessed on their practicability and solution approaches have been shown.

Many of these lead to the use of a voxel based direct representation of high complex geometries. The following assessment of this voxel approach has shown that it is feasibly to generate high complex geometries on basis of voxels.AcknowledgementsThe research leading to this results was carried out in the project OptiAMix – “Multi-target- optimized and continuously automated product development for additive manufacturing in the product development process”. This research and development project is funded by the German Federal Ministry of Education and Research (BMBF) within the program “Innovations for Tomorrow’s Production, Services, and Work” (funding number 02P158131) and managed by the Project Management Agency Karlsruhe (PTKA). The author is responsible for the contents of this publication.Literature[Ada15]ADAM, G. A.

COMPUTER MODEL FOR LASER PHOTOPOLYMERIZAnONLAWRANCE FLACH Department of Chemical and Materials EngineeringandRICHARD P. CHARTOFF Rapid Prototype Development Laboratory and Center for Basic and Applied Polymer ResearchUniversity of Dayton Dayton, Ohio, U.S.A.AbstractA computer model for a laser induced photopolymerization process has been established which simulates stereolithography. The model couples irradiation, chemical reaction, and heat transfer equations to provide insights into rate processes occurring in the volume element contacted by the laser beam. Quantities predicted include the spatial variation in conversion of monomer to polymer, depletion of photoinitiator, and local variations in temperature in and around the spot contacted by the laser.

Quantities predicted include the spatial variation in conversion of monomer to polymer, depletion of photoinitiator, and local variations in temperature in and around the spot contacted by the laser. This allows predictions to be made about the laser dwell time, depth penetration and uniformity of the photopolymer formed in the process.IntroductionOne aspect of the stereolithography process that has not received much attention during initial development of the overall procedure concerns the physical and chemical changes that occur in the small zone of monomer/polymer mixture directly exposed to laser irradiation, and the rates at which these changes occur.

The work described here addresses this aspect of the process.Mathematical models have been developed for the photopolymerization process which allow computation of light intensity, absorbed light, photoinitiator concentration, monomer concentration, and temperature profiles, as functions of time, for the exposed reaction zone and surrounding monomer/polymer mixture. The model discussed here·.deals with· the situation where the laser light spot remains stationary on the surface of the resin vat. A second model, one capable of handling the situation where the light spo!mOvesacross the surface of the resin vat in a series of small discrete steps, is being developed. Some preliminary results obtained from the moving spot model are presented to illustrate some of its capabilities.155Mathematical.•• ModelThe mathematical model for the stationary light spot situation was developed by considering a cylindrical region of ex.posedmaterial as shown in Figure 1.

For this situation, where the light spotdoes<not move with time, and the resin is spatb~lly homogeneous, cylindrical symmetry· can be assumed. This allows .light intensities,concenttations and temperatures to be functions of the two spatial coordinates (r and z) and time (t).The incident laser light is assumed to have a Gaussian intensity distributionI(r,z,t) ==I o e2 (..!) 2 Rat z==O, r~O , t~OThe beam radius, R, being the value of r at which light intensity drops to e-z of the peak: value.The decrease in light intensity with depth into the resin due to absorption by photoinitiatoris given byaI == -eSI azwith the light intensity at the surface (z=O) given by eqn.(l). The light absorbed by photoinitiator is thus given byIII.

(r , z, t)== eSI •Since photoinitiator is consumed by the photochemical initiation process, the rate ofchange of photoinitiator concentration can be calculated fromas == -~I atawith S(r,z,t) = So at t=O, r~, z~.Monomer is consumed by the polymerization reaction, so the rate of change of monomerconcentration can be calculated fromwith M(r,z,t) =Mo at t=O, r~, z~.156(1)(2)(3)(4)(5 )According to Odian [1], a rate expression of the following fonn can be used for aphotochemical polymerization process:cl>I ]0.5 a R;p = k;p M k e[If an expression of this fonn is valid then eqn.(5) becomes aM = -k M [ cI>.Ia ]0,5 .at;pk t;The temperature field in the vicinity of the exposed region is given by a transient heatconduction equation with heat source tenn (heat generated by polymerization) [2]withT(X,z,t) = To aT = 0 ax T = To T = Toat t=O, X:I?;O , Z:I?;Oat X=O, Z:I?;O,t:l?;Oat x .....oo, Z:I?;O,t:l?;Oat z .....oo, X:I?;O , t:l?;OaT = 0 azat z=O, X:I?;O,t:l?;OThe partial differential equations presented above together with the appropriate boundary It should be realized that this is a In addition, a number of assumptions areand initial conditions constitute the mathematical model.

model and is an approximation to the real situation. implicit in the model fonnulation as presented here. Briefly these are as follows:1. No flow of material (convection or diffusion) in any direction. 2. No optical effects (scattering, diffraction, refraction or reflection). 3. Heat generation due to heat of polymerization only. 4. Negligible heat loss from surface of monomer vat. 5. All physical and thennodynamic parameters constant. 6. Photoinitiator fragmentation products non-light-absorbing. When considering a mathematical model1. No flow of material (convection or diffusion) in any direction. 2. No optical effects (scattering, diffraction, refraction or reflection). 3. Heat generation due to heat of polymerization only. 4. Negligible heat loss from surface of monomer vat. 5. All physical and thennodynamic parameters constant. 6. Photoinitiator fragmentation products non-light-absorbing.

4. Negligible heat loss from surface of monomer vat. 5. All physical and thennodynamic parameters constant. 6. Photoinitiator fragmentation products non-light-absorbing. for157( 6)(7)(8 )transient heat conduction equation in three spatial dimensions, x, y, and z, is required. Motion of the laser light spot is achieved simply by changing the coordinates of the light spot center on the surface of the monomer vat.of.couplodpartial.•differential••equations of different types, the most complex of .which is the· parabolic, second order· heat conduction equation. In order to set up numerical solution. techniques for the set of equations, the infinity boundary conditions for the heat conduction equation were approximated by specifying the temperature at some fmite r and z values. It has been found that values of r=5R and z=2D are adequate for the situation used to test model operation (see Figure 2).

It has been found that values of r=5R and z=2D are adequate for the situation used to test model operation (see Figure 2). This approximation is only really valid for short times «0.5 seconds) as for longer times the boundary conditions start to influence the computed temperature values. Both Runge-Kuttaand Alternating Direction Implicit methodsThe.mathematicallllodel·thusconsists of a .systelll4] were used to numerically solve the equations.Results and DiscussionThe mathematical model presented above is general in that it can be used to model any resin system provided the appropriate physical, thermodynamic, and kinetic parameters are supplied by the user. The model was tested using physical and thermodynamic properties for a HDDA monomer, and a 15 mW laser light source. This information together with the dimensions of the exposed region, etc., Polymerization rate information for HDDA was obtained from a calorimetric study by Tryson and Shultz [5].

This information together with the dimensions of the exposed region, etc., Polymerization rate information for HDDA was obtained from a calorimetric study by Tryson and Shultz [5]. These data were used because they represented the most comprehensive available data set containing temperature dependent rate information. The data and kinetic parameters were derived experimentally by curve fitting to the 1/2...order kinetic expression of eqn.(6). Thus, in order to calculate consistent results it was necessary to use the 1/2-order rate expression in our treatment. It should be noted that mechanistically an expression that is 1st-order with respect to light intensitY is most appropriate for photopolymerization of multifunctional monomers [6]. The computer model is completely general, however, and will accommodate any rate expression with appropriate rate constants.is summarized in Table 1.Figures 3 and 4 are contour plots prepared from data obtained from the stationary spot model.

Figure 3 illustrates how conversion of monomer to polymer varies with position an exposure time of 0.3 seconds. The contours clearly illustrate how conical or "bullet-shaped" test pieces would be formed by short.;term exposure with a stationary laser beam. Figure 4 shows the corresponding temperature field in the vicinity of the exposed segment.Figures 5 and 6 represent some preliminary results obtained from the moving spot model currently under development. Figure 5 shows temperature profiles along the y-axis, and how these change with time as the light spot moves in the positive y...direction. It is interesting to note that for short times «0.3 seconds) the center of the laser spot precedes the peak in the temperature distribution, whereas for longer times (>0.4 seconds) the center of the spot lags behind the temperature peak. The fact that the resin is heated prior to exposure may be an important factor which could affect the nature of the polymer formed during the process.

The fact that the resin is heated prior to exposure may be an important factor which could affect the nature of the polymer formed during the process. Figure 6158shows the temperature contours on the surface of the monomer vat for an exposure time of 0.3 seconds.ConclusionA mathematical model for the stereolithography process has been developed and tested for a HDDA monomer reaction system and stationary laser light source. The model can be used to investigate process dynamics in slices of various thickness, with different laser light intensities, exposure times, and physical and chemical parameters. Results obtained allow prediction of laser dwell time, light depth penetration, and uniformity of polymer formed during the process. Model predictions are reasonable, but experimental verification has yet to be performed. A model for the moving spot situation is currently under development,and preliminary results obtained look promising.AcknowledgementsThe authors gratefully.

Due to the fact that rapid prototyping processes do not allow the direct manufacturing of high precision parts, it is very often necessary to measure a first partin order to modify the process parameters.After introducing the digitization means, basedona CMM machine and a plane laser sensor, the simulation environment will be presented as adapted for simulation and validation of 3D-laser scanning paths. The CAPP (Computer Aided Process Planning) system used for the automatic generation· of the laser scanning process will also be introduced. Keywords: scanning process simulation, inspection, optimization1 INTRODUCTIONcycle of new products needs time compression, from early design stages to industrial production. Rapidprototyping (Bernard & aI., 1998) is very often used to produce test parts representative of CAD model. The objective of digitalization consists, starting from a physical object, in acquiring a digital image of the parts on a point cloud form.

The objective of digitalization consists, starting from a physical object, in acquiring a digital image of the parts on a point cloud form. After a processing to make it coherent, this cloud is exploitable by CAD software for the surface reconstruction of the part, its manufacture or its measurement (Figure 1).surfacesproject, and due to industrial demands, the main objective is measurement of prototype parts or tools, in order to verify their conformity to the reference CAD model.

In this field of application, the surface (or volume) efficiency of the fITst step of the process "scanning modeling or technological applications (STL, CNC, ...)" (Figure 2) is one of the main important steps of control process because it decides.ofhow the part will be digitize, and mainly ofthe final result.point cloud treatmentThe challenge of the presented project isautomatically generate the measuring process of the partin order to be able to characterize the errors compared toreference model.This complete goal contains differentvalidate a scanning process for a plane process and the corresponding sensor paths.one is to be able to simulate and to to automatically generate this scanningcontext is defined as follow:the digitalization environment (CMM, sensor) is known and thechinematics parameters are defined, the CAD modelofthe part to be measured is the reference and is numerically defmed.

factCAD model of the part is known involves that this is not reverse engineering objective aI., 1998)· (Zhang &.al., 1995) but the goal is the inspection (or measurement) result and the,",VJ..JL""""""'J..J..","'J..J.. ofpoinfclol.ldswith the CAD model (figure 3).rapidly digital data from scanning processiin order to obtained' by digitizationreference CAD model andmain application fields of this application are rapid product development, and more especially ......."''''T·.'''.... applications. So, geometry of parts is very. complicated and mechanical data acquisition. In order to reduce.time consuming and to.improve point .... " " ". . rlI.c.rlI to use a plane laser sensor. One of the major improvements of the......... ,........ '"129complete inspection process using laser sensor is to automate off-line generation, simulation and validation ofthe laser sensor scanning process.Scanning strategyI Point clouds ICAD modeling,...

Afiera briefanalysis ofthestepspf scanning, th~proposed simulation and validation of 3D-laser sensor scannmgpr()c~.ss\Vil1 b~presented. Finally, some elements for the automatic generation of the scanning process (<letermj.Ilatjonofther~la,ted positions between part and sensor, trajectory definition) will also be introduced and argumented.Scanning strategyPoint cloud computingSurface modelingScanning process planningFigure 2. Scanning and modeling processCAD modelError mapPoints cloudsFigure 3. Control objective and environment2 SCANNING WORKCELL Laser sensors are materially unable to completely digitize a complex partin only one relative orientation between sensor and part. According to part complexity and according to sensor technical specifications, it remains areas where digitalization is not possible. These areas are the result of shadow or occultation phenomena depending.

These areas are the result of shadow or occultation phenomena depending. whether the surface part is seen by the camera but non-enlightened, or the surface indeed enlightened but unseen by.the camera.Figure 4. Coordirulte:measuring machineThe digitization cell is made with a coordinate measuring machine (CMM) GAMMA 1203, from DEA(figure 4) and a laser sensor from Kreon Industrie (Kreon Industrie, 1997). It is made of the 3D planeJaser sensor KLS 51 (Figure 5), an Electronic Control Dnit (ECD) and the "Kreon Reporter Plus" softWare. This environment has been detailed in (Davillerd, 1998).131Plan image: matrice CCDPlan laser2 CCD cameraslaser plane: triangulation on the laser lineFigure 5.

This environment has been detailed in (Davillerd, 1998).131Plan image: matrice CCDPlan laser2 CCD cameraslaser plane: triangulation on the laser lineFigure 5. KLS 51 laser plane sensorFurthermore, it can be noticed that thanks to the software and to the mechanical interface between theCMM and the sensor, the Kn50n acquisition system is able to manage 4 positions of sensor, around the axis, imperatively parallel with machine axes.Based on the active triangulation principle, the sensor uses plane laser technology and is equipped with two CCD (Charge Coupled Device) cameras, placed here and there of the laser plane.

of two cameras instead of only one allows attenuating the problems involved in the shadow or occultation phenomena on complex parts and increasing the number of points acquired with each digitization.Kreon acquisition system has a measurement field of trapezoidal form with a 45-mm depth.characteristic confers on the user facilities for the creation of trajectories by reducing the movements necessary to place the surface to be scanned in the measurement field.3 DIGITIZATION IMPROVEMENTS In this environment, a complete. analysis ofthe digitalization process has been proceeded to fmd the critical stages, which will be studied and improv~d..A flowshartdescribinK this process is detailed in (Davillerd, 1998). It has been. decomposed into three main steps: scanning process definition, scanning process execution, and data acquisition computing. In the following, each of these three steps will be detailed.

It has been. decomposed into three main steps: scanning process definition, scanning process execution, and data acquisition computing. In the following, each of these three steps will be detailed. 3.1.Scanning process defin.ition.This step consists in severalstages .like: digitization strategy definition, machine calibration, sensor calibration, software origin definition, part set-up, and strategy implementation. Only two of them are very critical: digitization strategy defmition and strategy implementation. stage. in which the user. defmes part orientations and sensor Digitization strategy definition is orientations that needed to digitize the part. Also, he has to define reference features to be able to readjust point clouds in a next stage. These referenc.e features can consist in part features or in features added on the part or in a rotation axis.derme sensor trajectories.

These referenc.e features can consist in part features or in features added on the part or in a rotation axis.derme sensor trajectories. So he programs the coordinates of linear movements and indicates to the system the moment along with it has to digitize (Figure 7). He also chooses the acquisition step. In this way, he also implicitly defmes the recovery (number of point that will be digitized in more than one position, and that will have to be filtered).Strategy implementation is very time consuming. Indeed,the user has here132Laser3.2..Acquisition execution.Starting from the trajectory file defmed by the user, the ECD pilots the sensor in the machine reference system and Kreon Reporter Plus software records the acquired points. Everything is automatic and has not to be improved. 3.3.Data acquisition computing. It consists in two steps. First, point cloud filtering is necessary because the acquisition generates an important quantity of points.

3.3.Data acquisition computing. It consists in two steps. First, point cloud filtering is necessary because the acquisition generates an important quantity of points. Consequently a simplification is necessary and consists in grid application. But the needed quantity and quality of points essentially depends on the using application: for examplemachining, .sllrfaceireconstruction or. imaging.And finally, the readjusting stage is very important because it conditions the final results. Each point cloud is defined in a particular frame then all frames have to be regrouped in a unique one. This operation can be made manually or by using matrices. Then, after the part is digitized according to the defined strategy, the whole numerical information ofthe part is defined (figure 6).Physical objectNumerical modelFigure 6. The part on its fixture and the fmal point cloud.AND VALIDATIONa "step by step" process.

The part on its fixture and the fmal point cloud.AND VALIDATIONa "step by step" process. This means that the CMM moves, then stops its movements, the the CMM moves again. Thus, each acquisition is a discrete event. Let us detailsensor acquires data this event forthe 3D-laser-scanning sensor KLS 51.The fITst stage is the part illumination by the laser. A plane laser lights the part up. A beam coming from a laser diode creates this plane. This beam goes through a cylindrical lens in order to create the plane. This plane is about 0.3 mm thick in the middle ofthe measurement field.The second stage is the scene acquisition. Each camera takes a gray level image of the scene. EachCCD camera acquires the laser trace on the part and its thickness.The third and last stage is the image computing. Each image is processed in order to generate 3D order to keep .interesting information. This treatment trapezoidal sensor measurement field.

Each image is processed in order to generate 3D order to keep .interesting information. This treatment trapezoidal sensor measurement field. A numerical treatment called sub-pixel algorithm.is then used to reduce thickness to one point. Triangulation calculation allows the system to extract the three dimensional coordinates of300 points on the traceJength (Kreonlndustrie, 1997).coordinates of acquired points. First, image is cutsimulation is tovisualize.CMMlllovelIlentsand.eachacquisition of the sensor. So, it has. been.. necessary to model the digitizationcell.in . robotic CAD· environment (CIMSTATION INSPECTION from Silma). Several applications exist (Chedm~il & aI., 1998): spot welding, arc welding, painting and projection are available. Collision between the sensor and the digitization environment can be detected as well.analysis of this environment shows that sensor actions could be similar to paint deposition.

Collision between the sensor and the digitization environment can be detected as well.analysis of this environment shows that sensor actions could be similar to paint deposition. beam action is the illumination of part surface and it could be compared with.........u ........., 1998). Indeed,133a paint deposition on the lighting area. In the same way, camera action is the part surface observation and it could be compared with a paint deposition on the viewed area. In this way, the plane laser sensor work has to be adapted in the chosen robotic CAD software from the paint deposition simulation.So, a solid model of the laser sensor has been adapted on the head of the machine in CIMSTATION ROBOTICS environment. Attachment, 4-positions interface and sensor surface models· are reconstructed from IGESCAD files with CIMSTATION functions (Figure 7).Figure 7.

Attachment, 4-positions interface and sensor surface models· are reconstructed from IGESCAD files with CIMSTATION functions (Figure 7).Figure 7. Models of machine, sensor and simulation features.Original models have been proposed for the •laser and each camera, adapted from paint plumes,which have specialpaint deposition function. Laser plume model has been created from a cone plume.. The measurement field restriction treatment has beenintegrated inthis plume; i.e. laser plume paint deposition will only be effective in the measurement field area.· Eachca,mera plume model has been created from a cone plume too. Then the intersection of the area painted by the laserplllme and the area painted by either camera plume corresponds to the digitized points.

Then the intersection of the area painted by the laserplllme and the area painted by either camera plume corresponds to the digitized points. Sub-pixel algorithm allows a compromise between laser width, mesh size of the CAD model and a realistic time to execute calculations in order to keep an interactive exploitation ofthe results.In conclusion, digitization .cell and sensor have been modeled (Davillerd, 1999). It is now possible for each particular 3D part to simulate digitization trajectories that shouldhaveheen previQusl)'icrea,ted. Based on these elements, off-line simula,tion allows· .testing and verifying digitization. tra,jectories efficiency without immobilizing the digitization cell. More over, it is also possible to valida,te these trajectories by the analysis of the valid points (enlightened and seen by one or two cameras) obtained using the simulation module.

More over, it is also possible to valida,te these trajectories by the analysis of the valid points (enlightened and seen by one or two cameras) obtained using the simulation module. The analysis algorithm is under validation.5 COMPUTER-AIDED PROCESS PLANNING FOR AUTOMATIC GENERATION OF 3DDIGITIZING PROCESS FOR LASER SENSORSIn this paragraph, an analysis is proposed in order to generate an automatic scanning process. As mentioned earlier, the objective is control. In this field, two configurations are possible. At frrst,the CAD model of the part is known and it has never been scanned. It is almostnecessary to generate the process· to digitize the. part. The ·second hypothesis is that the part has already been scanned manually andthe result is not satisfactory (empty areas or too much redundant points). In these cases, the proposed algorithm has todefme anew efficientscanning process. Infact, we thinkthat two stages are necessary to generate an automatic scanning process.

In these cases, the proposed algorithm has todefme anew efficientscanning process. Infact, we thinkthat two stages are necessary to generate an automatic scanning process. Theflrsfstage is the determination of the different orientations between the sensor and the part, and the second stage is the defmition of the trajectories,for each orientat1on,which allow digitizing the maximum part surface. In the following, we suggest different concepts to solVe these two stages.5.1 Definition of the relative orientations between the sensor and the part. For the flrststage, we need to know whether the surfacemeasurement plan and whether one of the two cameras sees the entire or laser on the surface element S;. These conditions refer to a visibility concept. It haswith different application fields, 1989) and inspection with computer vision (Trucco & aI., 1997). Into IIl~charricalprobe (Spyridi & machining and ·inspection with a mechanical probe field,.

It iS~axtitjone~intocate~ories.andpapersto path generation classified according to the topics they cover. In t~esame.way,Chedmail(1998)describes the·differentmethods ofpath planning in robotics. The goalwill be,a.fteradaptationand transformation of these methods, to optimize the trajectories for each related position<between the part andthe sensor.1355.3 Constraints for the automatic scanning process generation.Different solutions have been introduced to generate an automatic scanning ........r."".:."c-C'several constraints that· must be taken into account during the algorithm development.One of these constraints is the. recovery, which is not. mastered by the user when he defines the process. When the system. defines a trajectory, it takes account to the recovery othef'\¥ise there)lllaybe redundancy or missing of points.

mastered by the user when he defines the process. When the system. defines a trajectory, it takes account to the recovery othef'\¥ise there)lllaybe redundancy or missing of points. This recovery notion(figure 10) is difficult t0qlIlsterbecauseofthe laserof the measurement zone, the trace projected by thelaser is slllallest than>whel1thepart surface is located at the bottom of the zone.Figure 1O.Path recovery.Concerning the measurement zone, the acquisition is possible in the all weight of the zone, i.e., the sensor can move in the same z-coordinate; the part surface can vary in z, and stay in the measurement zone. The difficulty in this case is to model the "tool". For example, in machining, the tool.works with its end. Italso. true. with a touch probe. It is more complicated to specify the active part ofthe laser plane. Another constraint is the angle· between the laser beam and the part surface.

Italso. true. with a touch probe. It is more complicated to specify the active part ofthe laser plane. Another constraint is the angle· between the laser beam and the part surface. Even if the algorithm determines· a visibility, it must be verified that the angle does not keep away from the normal of the surface.6 CONCLUSION AND OUTLINES Control of complex part needs to be improved in term of scanning efficiency. This contribution isrelated to "simulation validation - CAPP for generation of 3D-laser scanning process". At this time, the. fITst step has been implemented and ·the second one is under development. The third one will be· achieved in October 1999, based on spherical geometry.The final complete system will contribute to optimize the development of new products and toaccelerate the control phase.REFERENCESA., Taillandier, G.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCOMPUTER-AIDED PROCESS PLANNING FOR WIRE ARC DIRECTED ENERGY DEPOSITIONYi Xiong1, Audelia Gumarus Dharmawan1, Yunlong Tang1,Gim Song Soh1,2, David William Rosen1,2,31Digital Manufacturing and Design Centre, Singapore University of Technology and Design, Singapore, 4873722Engineering Product Development, Singapore University of Technology and Design, Singapore 4873723The G. W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology Atlanta, GA 30332AbstractWire arc directed energy deposition provides a rapid and cost-effective method for fabricating low-to-medium complexity and medium-to-large size metal parts.

However, the complex nonequilibrium phase transformations, inherent to this process, make it a challenging task to produce consistent and high-quality parts, especially for parts with materials or geometries that have not been manufactured before. This study outlines a holistic and data-centric computer- aided process planning framework utilizing a knowledge base to assist engineers in selecting optimal process parameters that reduce dimensional deviations, and therefore to obtain near- net-shape parts using directed energy deposition only. The knowledge base has a data- knowledge-service architecture and is proposed to synthesize information from various sources, e.g., characterization tests. Based on these collected data, several knowledge representations, including database, metamodels, and planning rules, are constructed to support decision- making in the process planning.

Based on these collected data, several knowledge representations, including database, metamodels, and planning rules, are constructed to support decision- making in the process planning. The proposed framework is demonstrated in the fabrication of components from industrial applications.Keywords: Process Planning, Directed Energy Deposition, Knowledge-Based Engineering, Design for Additive Manufacturing1. INTRODUCTIONDirected energy deposition (DED) refers to a category of additive manufacturing (AM) processes that utilizes a high energy source to melt the feedstock material, which comes in either wire or powder form, and patterns this energy source with motions provided by a computer numeric control (CNC) machine or robot to infill enclosed regions for creating a single layer of a three-dimensional (3D) part.

The DED process is suitable for fabricating low- to-medium complexity and medium-to-large size metal parts, in which other metal AM processes such as powder bed fusion (PBF) cannot be used [1,2]. Due to its high deposition rate and low setup cost, DED offers an efficient and effective method for applications that require large near-net-shape parts at a short lead time. Hence, it has been utilized in the maritime, aerospace, and automotive industries with applications including marine propellers [3], winglets [4], excavator arms [5]. In addition to creating new parts, DED can also be used for on-site remanufacturing applications due to its portability and large work volume when robots are used as motion devices.1094Although wire arc DED has many favorable characteristics, several challenges should be solved to make this manufacturing method easier to operate and more reliable for industrial production.

The most difficult challenge is that the physical and chemical processes within manufacturing, e.g., nonequilibrium phase transformations are complex and not well understood. The lack of understanding of these underlying processes has greatly hindered the development of optimal process planning strategies for fabricating consistent and high-quality parts. To address this issue, many studies have been conducted to establish process-structure- property relationships and to utilize these relationships for exploring optimal process planning strategies.However, these studies are often ad hoc that are either focus on specific materials or specific geometry features such as thinned-walls. Additionally, these data, models, and knowledge acquired are not correlated and formalized for reuse.

Additionally, these data, models, and knowledge acquired are not correlated and formalized for reuse. More importantly, current computer-aided process planning tools are mainly based on frameworks that were developed for either less complicated AM processes such as fusion deposition modeling or subtractive manufacturing such as milling. Hence, these CAPP tools often do not allow adaptive process planning for complex features. Above limitations in existing methods and tools necessitate a computer- aided process planning framework for wire-arc DED that considers different materials and geometry features, allows capitalization of existing knowledge, and evolves with times.To support this view, this study outlines a knowledge-based computer-aided process planning framework to assist operations engineers in selecting optimal process settings for obtaining high-quality near-net-shape parts through the wire arc DED process.

More specifically, the knowledge base about the wire arc DED process is constructed with a data-knowledge-service architecture to capture, represent, and reuse domain knowledge. The rest of the paper is organized as follows. Section 2 reviews related works in this field. The data-knowledge-service based computer-aided process planning framework is illustrated in Section 3. Section 4 presents the software tool implementation of the proposed process planning framework. Finally, conclusions are drawn in Section 5.2. RELATED WORKThis section reviews relevant work on two research topics: process planning for wire arc DED and AM knowledge management. This will serve as the basis for the research presented in this paper.2.1. Research in process planning for wire arc DEDProcess planning is a significant step in the AM workflow that links design and manufacture.

Research in process planning for wire arc DEDProcess planning is a significant step in the AM workflow that links design and manufacture. It converts geometric models together with their functional requirements into process plans that are to be executed by AM machines. According to [6], the conversion is supported by two types of activities: decision-making processes and algorithmic operations. The decision-making process searches for compromise process settings that satisfy multiple goals. Then, these settings are converted into process plans through algorithmic operations. This general categorization is also valid for the wire arc DED process, and we review literature in these two categories.Decision-making for wire arc DED is a challenging task because the underlying mechanisms are complex and not well understood.

Hence, research efforts have been made to establish relationships between process variables and properties of fabricated parts such as geometries[7], surface roughness [8], and mechanical properties [9]. Both experimental and theoretical tools have been used to reveal the physical mechanism. For instance, thermal imaging techniques were utilized in [10] for obtained the surface temperature field of deposited parts to study the1095effect of interlayer cooling time on the surface quality and part geometries. In addition, the internal flow behavior of melt pools was studied [11] through a high-energy synchrotron radiation experiment. Meanwhile, analytical and numerical models have been developed to predict bead geometries [12] and residual stresses [13], respectively.

Meanwhile, analytical and numerical models have been developed to predict bead geometries [12] and residual stresses [13], respectively. Finally, these process- property relationships were represented in various forms such as response surface methodology [14] and neural networks [15] to support engineers to define process windows and settings for the arc DED process.Meanwhile, various process planning algorithms have been developed for the wire arc DED process. For instance, reference [16] developed a medial axis transformation (MAT) based toolpath generation algorithm to increase the production speed and to improve the geometry accuracy. This algorithm was further developed in [17] [18] with more robust solutions such as the level-set method to generate toolpaths with variable bead width. In addition, sequential path planning methodology based on a water-pouring rule has been investigated in [19] to enable real-time hatching with a feedback loop.

In addition, sequential path planning methodology based on a water-pouring rule has been investigated in [19] to enable real-time hatching with a feedback loop. Some recent research efforts [20,21] focused on feature/modular-based algorithms in which specific toolpaths are generated for various manufacturing features such as free end walls, t-crossings, and direct-crossings, which are segmented from a design part.2.2. Research in AM knowledge managementPrevious research has addressed the need for knowledge management in AM that allows information storing and information reuse to support decision-making. AM knowledge was defined to include design knowledge, manufacturing knowledge, and other knowledge within the AM value chain. Methods from knowledge-based engineering (KBE) have been used as primary tools. Ontologies for AM have been proposed for knowledge formation and reasoning.

Methods from knowledge-based engineering (KBE) have been used as primary tools. Ontologies for AM have been proposed for knowledge formation and reasoning. For instance, a computational ontology has been proposed in [22] for representing and reasoning over AM knowledge and data. A design for AM ontology has been developed in [23] primarily for knowledge documentation using the formal web ontology language (OWL)/ resource description framework (RDF). This ontology was used for manufacturability analysis. Also, the innovative capabilities of additive manufacturing (ICAM) ontology was proposed in [24] for innovative design ideation.Other research focused on the development of hierarchical models for knowledge management. Reference [25] presented a hierarchical object-oriented model with a seven-layer data organization that covers steps from the manufacturing planning until post-production.

Reference [25] presented a hierarchical object-oriented model with a seven-layer data organization that covers steps from the manufacturing planning until post-production. In addition, a four-tier framework based on a closed-loop data-information-knowledge- application framework was proposed in [26] for self-improving additive manufacturing knowledge manufacturing.Research gaps are evident with the above studies on AM knowledge management.1) The existing AM knowledge bases were targeted at general AM processes whereby unique characteristics and requirements of AM knowledge base for a specific AM process, i.e., wire arc DED were not explored and identified. 2) The method to utilize AM knowledge base for CAPP was not studied. 3) Additional research is needed on the user interface characteristics of knowledge base systems, such as ease of use, support to designers as services, and classification of users.

3) Additional research is needed on the user interface characteristics of knowledge base systems, such as ease of use, support to designers as services, and classification of users. In this paper, a data-knowledge-service architecture is proposed for capitalization and reuse of manufacturing knowledge about the wire-arc DED process.3. KNOWLEDGE BASED COMPUTER AIDED PROCESS PLANNINGThis section first provides an overview of the proposed data-knowledge-service architecture for wire-arc DED. Each level of the knowledge base is explained in detail at each subsection.10963.1.Data-knowledge-service architectureThe proposed knowledge-based computer-aided process planning framework consists of three layers: data, knowledge, and service. As shown in Figure 1, these three layers are organized in a hierarchical structure that reflects the abstraction process from data to service.

As shown in Figure 1, these three layers are organized in a hierarchical structure that reflects the abstraction process from data to service. The basic layer includes data obtained from three sources: process characterization, in-process monitoring, and post-process measurement. Based on the data layer, knowledge is stored in various formats, such as metamodels, database, and planning rules. The above knowledge is then used as a basis for two types of services: 1) supporting operations engineers in selecting process settings; and 2) communicating with product designers regarding the potential design freedom in the domain of process control. The first type of service can be viewed as an extension of existing process planning to consider complex AM processes.

The first type of service can be viewed as an extension of existing process planning to consider complex AM processes. The second type of service is a unique service in the context of design for additive manufacturing which emphasizes the possibility to explore new design freedom.The use of the data-knowledge-service architecture brings several advantages to move wire arc DED from the R&D arena towards the production line. First, the bottom layer that is data- driven stores information in a central repository and fills the increase needs of part quality assurance and qualification. Second, the middle layer that is knowledge-based offers an integrated method to capture and represent knowledge which moves process planning from an information intensive basis towards a knowledge intensive foundation.

Finally, the top layer that is service-oriented provides manufacturing bureaus a way to formalize their know-how for guaranteeing their competences.Product Operations engineer designerData-Knowledge-Service ArchitectureRole:Service LayerInterface to different stakeholders Minimizes the requirement of specific knowledge to the end userRole:Knowledge LayerTo capture and represent knowledge Offers a knowledge intensive foundationRole:Data LayerCentral repository for storing and managing dataFigure 1 A knowledge base with a data-knowledge-service architecture for computer aided process planning3.2.Data layerWire arc DED utilizes a complete digital workflow in design and manufacture. Various data are generated at different steps of the workflow. The primary function of the data layer is to store these data in a structured manner for retrieving.

Various data are generated at different steps of the workflow. The primary function of the data layer is to store these data in a structured manner for retrieving. Based on their sources, collected data are categorized into three groups: process characterization, in-process monitoring, and post- process measurement. Heterogeneous data are stored in each category that includes various formats, e.g., CAD models, point clouds, and images.1097For wire arc DED, process characterizations are often conducted for new materials and new geometric features, as these variations significantly affect the process. Hence, process settings must be adjusted accordingly based on the newly established relationship between process settings and performance of fabricated parts. Typical process characterization tests include: single-bead single-layer build; multi-bead-multi-layer-build; singe-bead multi-layer-build; and builds for specific geometry features such as overhangs.

As summarized in Figure 2, each test is designed to gain an understanding of the process from a certain aspect. For instance, the single-bead single-layer build is conducted to define a process window from where feasible process settings are selected. However, for bulk parts and thin-walled structure, their distinct conditions for the phase transformations require specific investigations. In this context, multi- bead multi-layer and single-bead multi-layer builds are fabricated and the part geometries such as the layer height and wall thickness are measured.

In this context, multi- bead multi-layer and single-bead multi-layer builds are fabricated and the part geometries such as the layer height and wall thickness are measured. For specific geometry features, the process settings should be adjusted according to feature characteristics such as hole sizes and overhang angles to obtain satisfactory performance of fabricated parts.Single-bead single-layer buildBead geometriesProcess parameter (PP) window/ Step-over ratio2 Multi-bead multi-layer buildLayer heights/ Part geometriesOptimized PP for bulk structures3Single-bead multi-layer buildLayer heights/ Effective wall thicknessOptimized PP for thin structures4Overhang structureLayer heights/ Part geometriesDataset I metamodel1.

Single-bead single-layer build2.Single-bead multi-layer build3.Multi-bead multi-layer build4.0verhang structureFigure 2 Process characterization dataNowadays, more and more sensors are installed in wire arc DED systems for in-process monitoring. On the one hand, condition monitoring sensors, e.g., temperature sensors, high- speed cameras, and photodiodes offer data about the melt-pool, thermal distribution, and thermal history for facilitating understanding of underlying mechanisms that are hard to model and predict. On the other hand, online non-destructive testing (NDT) methods, e.g., infrared thermography and laser thermography provide data that include layer defects information for the increasing demands of part qualification. In-process monitoring data are acquired at a high sampling rate and often stored as images.

In-process monitoring data are acquired at a high sampling rate and often stored as images. This requires specific data infrastructure and data processing methods, such as data compression and feature extraction.Other data comes from the quality check for fabricated parts. These data contain information about the achieved performance of fabricated parts such as mechanical properties, microstructure, and part geometries. These data provide feedback to compare against the desired performance and guide further corrections. For instance, since the part fabricated by wire arc DED is near net-shaped, its geometry is required for further post-processing such as1098machining.

For instance, since the part fabricated by wire arc DED is near net-shaped, its geometry is required for further post-processing such as1098machining. The fabricated parts are often scanned through a 3D scanner and the outcomes are point clouds, which are further processed and converted into geometric part models for downstream machining.3.3.Knowledge layerThe knowledge layer allows capturing and representing information stored in the data layer and supports the service layer for decision making by utilizing knowledge. Because of the heterogeneity of data, knowledge is represented in various formats, such as metamodels, database, and planning rules.Metamodels are often used for describing knowledge that has an implicit form. Examples of these models are response surface models, radial basis functions, and Kriging models. These models are often constructed based on datasets obtained from either experiments or simulations.

Examples of these models are response surface models, radial basis functions, and Kriging models. These models are often constructed based on datasets obtained from either experiments or simulations. For instance, two response surface models for describing relationships between bead geometries and process variables are constructed as shown in Figure 3. The inputs of these models are often continuous variables that allow selections within a value range. For discrete variables that have finite values, a database can be used to store the relationship between inputs and output. For example, wire arc DED utilizes different sheltering gases for fabricating different metals and alloys. For fabricating stainless steels, 95% Argon and 5% CO2 is used as the sheltering gas, whereby for fabricating nickel aluminum bronze, pure Argon is used.

For fabricating stainless steels, 95% Argon and 5% CO2 is used as the sheltering gas, whereby for fabricating nickel aluminum bronze, pure Argon is used. Additionally, planning rules are used to store heuristic knowledge including know-how accumulated during fabrications and other knowledge that has an explicit form. For instance, for the tapper part shown in Figure 3, an outside-in scanning sequence will lead to a failed print, while an inside-out sequence will likely lead to success. It is assumed this is caused by the heat accumulation that is trapped within the part center. However, a complex heat transfer and thermal mechanical model is required to give a complete understanding of the physical phenomena.

However, a complex heat transfer and thermal mechanical model is required to give a complete understanding of the physical phenomena. These types of generally applicable heuristic knowledge are therefore stored as planning rules for reusing.Knowledge TypeMetamodelDatabasePlanning RuleKriging Model: Process-Bead Geometry RelationshipIf Part=Solid Then, Inside-Out SequenceSwitch: MaterialExample:Case= Stainless Steel Sheltering Gas= Ar(95%) and C02(5%)Outside-InCase= nickel aluminium bronze (NAB) Sheltering Gas= ArInside-OutEndFigure 3 Different knowledge types: metamodels, database, and planning rules3.4.Service layerThe concept of the service layer is introduced to the knowledge base for supporting end-users while addressing issues identified in Section 2.2. On the one hand, it is evident that the end-1099user will still play a pivotal role in process planning before the arrival of fully automated process planning tools.

On the one hand, it is evident that the end-1099user will still play a pivotal role in process planning before the arrival of fully automated process planning tools. Therefore, it is important to consider how to support the end-user in decision-making through services. In an ideal scenario, the designer does not need to fully understand the complex underlying relationships at lower levels for solving the decision problem. On the other hand, providing service is an important mechanism for manufacturing bureaus to capitalize on their know-how and to stay competitive. Given the fact that AM hardware is becoming more standardized and has fewer differences, the advantage of providing services to fabricate high-quality optimal products is more apparent.

Given the fact that AM hardware is becoming more standardized and has fewer differences, the advantage of providing services to fabricate high-quality optimal products is more apparent. Services should be provided to end-users including product designers and operations engineers for analyzing the process capabilities and constraints, for visualizing the consequence of different decisions, and for understanding the trade-off among different objectives.To illustrate the concept of service, we provide an example on selecting wire feed rate and scan speed for obtaining desired bead geometries. As shown in Figure 4, a design space exploration tool is constructed to link its performance space and design space. Once the designer adjusts the target bead geometries in the performance space, corresponding feasible design space is displayed accordingly. Since there are two design targets, i.e., target height and width, design space that satisfy each target is shown in Figure 4 (b) and (c).

Since there are two design targets, i.e., target height and width, design space that satisfy each target is shown in Figure 4 (b) and (c). The feasible design space that satisfies both goals is obtained by overlapping feasible design spaces for each goal. The above service is supported by surrogate models, i.e., Gaussian Process Regression models and Bayesian Network Classifiers, and single-layer single-bead characterization data. More details can be found in our earlier work [27].(a)Performance Space: Bead Geometry(b) Designs satisfy height targets2.5E .s :E a, 2 ·;;; ,, I ., "'CDc ~ E Q) .; ,, ll:'. 4 ., ., u. ~ ~1.545 Bead Width [mm)6911Scan Speed [mm/s)13(c)Design satisfy w idth targets(d )Design satisfy both targetsc ~ .s .!!! "' ll:'. 4 al ., u. ~ ~c ~ .s .!!! "' ll:'. 4 al ., u. ~ ~9111391113Scan Speed (mm/s]Scan Speed (mm/s)Figure 4 Design space exploration between process variables and bead geometries4.

4 al ., u. ~ ~c ~ .s .!!! "' ll:'. 4 al ., u. ~ ~9111391113Scan Speed (mm/s]Scan Speed (mm/s)Figure 4 Design space exploration between process variables and bead geometries4. EXAMPLE: SOFTWARE TOOL IMPLEMENTATION FOR OVERHANG FABRICATIONA software tool that embodies the proposed data-knowledge-service framework was implemented. The current version is developed for fabricating parts with overhang features,1100which is a common but challenging feature for the wire arc DED process. As shown in Figure 5, the user interface includes various buttons for users to interact with.

As shown in Figure 5, the user interface includes various buttons for users to interact with. Each section of the user interface is explained briefly as follows.CAPP for Wire Arc-OEDServiceOverhang Angle CheckToofpath GenerationAd8ptlve Process Parameter EstlflllltlonModelProcess Rules (Overhang)[Melomodel (Overhang)DataDotaset(Stainless Steel 316L)Dataset(Maraging SteeQFigure 5 The CAPP tool with a data-knowledge-service framework(cid:120) DataFor overhang features, characterization tests were carried out on a benchmark part that has various overhang angles (from 0 to 50 degree). Different process settings were used to fabricate these parts with stainless steel 316 L and the geometries of fabricated parts were measured using a laser scanner. In addition, the melt pool image and base plate temperature can also be measured.

In addition, the melt pool image and base plate temperature can also be measured. These data were stored within a dataset and can be updated for other materials such as maraging steels.(cid:120) KnowledgeBoth process rules and meta models were constructed in the knowledge layer. In the characterization tests, the maximum overhang angle using a flat deposition strategy was found to be 40 degrees. This is formalized as a design rule: if the overhang angle is larger than 40 degree, then the feature is not printable. In addition, a response surface was built based on a meta model to represent the relationship among the process setting, overhang angle, and resulting feature geometry.(cid:120) ServiceThree services are provided for designers in fabricating parts with overhang features. First, in the early design stage, the designer checks the manufacturability of their part designs based on the knowledge acquired about the capability of the wire-arc DED.

First, in the early design stage, the designer checks the manufacturability of their part designs based on the knowledge acquired about the capability of the wire-arc DED. Based on the feedback from the service, the designer makes necessary changes to the part design and iterates this process until the part fully meets the manufacturability requirement. Second, the toolpath generation service automatically hatches the layer slice based on the configured settings. Then, the knowledge regarding the design feature, process settings, and part performance is used to estimate the optimal process settings to guarantee a constant layer height over different overhang angles.1101The software tool developed is used in the process planning stage within the AM workflow to manufacture part designs with variable overhang angles as shown in Figure 6.

As explained earlier, the CAPP tool provides service, i.e., overhang manufacturability check to communicate with product designers about design constraints and freedom. If the part design contains features that are not manufacturable, it will be returned to the previous stage of the AM workflow, i.e., the product design stage. Once the manufacturability has been checked, operations engineers can use two services, i.e., toolpath generation and adaptive process parameter estimation to obtain an adaptive process plan for fabricating the part. In addition, both in-process monitoring data and post-process measurement data are fed back to the knowledge base to update and improve the service provided.

In addition, both in-process monitoring data and post-process measurement data are fed back to the knowledge base to update and improve the service provided. This allows the knowledge base to evolve with new data.Start [Initial Part Design]Part Redesign/ Part OptimizationOverhangManufacturability 111 ►Check ServiceDes igner., ..Operations EngineerToolpathPlanning Service 111 ►Process Param eter Estimation Seri\Ce111 ►Toolpath GenerationProcess Parameter E stimationProcess Characterizatio Data-----· .In-Process Monitoring DataW ire Arc OED ManfaucturingPost Process Measurement DataInspection & Quality ControlFigure 6 The proposed CAPP tool to support end-users in the manufacture of a part with overhang features.As shown in Figure 7 (a)-(b), a thin-walled part design with overhang features was fabricated with the CAPP tool developed. The 3D scanned model of fabricated model is shown in Figure 7 (c).

The 3D scanned model of fabricated model is shown in Figure 7 (c). It indicates that the fabricated part (in green color) is larger than the than the part design (in red color) with marginal volumes that allow post-processing, e.g., milling. To demonstrate1102the improvement achieved with the proposed CAPP tool, a reference part fabricated using a constant process setting is shown in Figure 7 (d). The fabricated part significantly deviates from the design part geometry.(a)(b)(c)Figure 7 (a) CAD model; (b) part fabricated with the CAPP tool; (c)3D scans of part manufactured with the CAPP tool; (d) 3D scans of part manufactured without the CAPP tool. Both scans are marked in green and the original design is in red5. CONCLUSIONSThis study outlined a holistic and data-centric computer-aided process planning framework utilizing a knowledge base for the wire arc DED process.

CONCLUSIONSThis study outlined a holistic and data-centric computer-aided process planning framework utilizing a knowledge base for the wire arc DED process. The knowledge base was constructed based on a data-knowledge-service architecture which can evolve with new data. In the context of design for AM, this CAPP framework communicates with both up and down streams of the AM workflow: 1) inform designers about freedom and constraints on product design; and 2) assists operations engineers to search optimal process settings for manufacture.This method brings several advantages. First, the use of the knowledge base offers manufacturing bureaus a way to formalize their know-how for guaranteeing their competences.

In addition, the modularized layer structure eases the decision-making for operations engineers in process planning as there is no need to understand complex physical phenomena within the manufacturing process.Future work will focus on examining data mining and machine learning methods that can transfer existing knowledge that was gained from fabricated parts onto process planning for parts with materials or geometries that have not been manufactured before.ACKNOWLEDGEMENTSThe authors acknowledge support from the Digital Manufacturing and Design (DManD) research center at the Singapore University of Technology and Design supported by the Singapore National Research Foundation.1103REFERENCE[1]Ding, D., Pan, Z., Cuiuri, D., and Li, H., 2015, “Wire-Feed Additive Manufacturing of Metal Components: Technologies, Developments and Future Interests,” Int. J. Adv. Manuf. Technol., 81(1–4), pp. 465–481.[2]Cunningham, C. R., Flynn, J.

The process involves spreading the powder in thin layers and then selective binding of the powder using a technology similar to ink-jet ptinting. Layers are added sequentially until a part is completed. 3DP has been used to make complex-shaped components from several monolithic matetials, including cOlnponents for use in structural applications. This paper focuses, however, on the ability to control microstlucture and local composition by 3DP. We envision cases where cOlnputer delived-microstructlu'es can be created by approptiate control of the ptinting paralneters. Thus, one can build components with the desired microstlucture independent of the complexity of the desired shape. Examples for both structural matetials and biomedical devices are discussed.IntroductionThree Dimensional Ptinting (3DP) is a solid free-form fabtication (SFF) Inethod used to create components directly from CAD representations and has been described in several publications[l-3].

Btiefly, 3DP creates Pal·ts by a layered ptinting process. The infonnation for each layer is obtained by applying a slicing algotithln to the cOlnputer Inodel of the part. An individual two-dimensional layer is created by adding a layer of powder to the top of a piston and cylinder containing a powder bed and the part being fabticated. The new powder layer is selectively joined where the part is to be fonned by "ink-jet" ptinting of a binder Inatetial. The piston, powder bed and part al'e lowered and a new layer of powder is spread out and selectively joined. The layering process is repeated until the part is completely ptinted. Relnoval of the unbound powder reveals the fabticated part. The process has been primalily applied to cermnic Inolds for Inetal casting. More recently, however, it has been used to make metal parts directly[4], structural ceralnic parts[5], and polymer pal'tS[6].1813DP is one of the most flexible SFF technologies.

More recently, however, it has been used to make metal parts directly[4], structural ceralnic parts[5], and polymer pal'tS[6].1813DP is one of the most flexible SFF technologies. The process can create parts of any geometry, including intelnal volumes (as long as there is a hole for the loose powder to escape). The support gained frOln the powder bed means that overhangs, undercuts, and intelnal volumes can be created. 3D Plinting can fonn any Inatelial that can be obtained as a powder. Further, because different matelials can be dispensed by different printheads, 3D plinting can exercise control over local matelial composition. Matelial can be deposited as particulate matter in a liquid vehicle, as dissolved Inatter in a liquid canier, or as molten matter.

Matelial can be deposited as particulate matter in a liquid vehicle, as dissolved Inatter in a liquid canier, or as molten matter. The proper placelnent of droplets can be used to create sUlfaces of controlled texture and to control the internal microstructure of the printed palt.The ability to control local cOlnposition is a unique feature of 3DP and is the major focus of this mticle. This type of control pennits fablication of Inatelials with computer derived microstructures or These m'e components in which their Inicrostructure is designed on a cOlnputer and built via the 3DP process. 3DP can selectively deposit matter within the structure of a cOlnponent so that composition can vary frOln point to point. The Inacroscopic shape of the cOlnponent can, however, be specified cOlnpletely independently. Thus, we envision components where both the macrostructure and Inicrostlucture are designed by cOlnputer and constructed by 3DP.

Thus, we envision components where both the macrostructure and Inicrostlucture are designed by cOlnputer and constructed by 3DP. Potential applications of such a technology are nmnerous, such as cOlnponents with anisotropic mechanical properties, Inicroengineered porosity, or constructing cOlnposite multilayer modules for electronic packaging. Conventional powder fonning technologies Thus, cannot provide demonstration of this approach will be a quantmn leap beyond CUITent Inaterial fablication and will create a technology that is not unlike the control that photolithography provided the electronics industry.spatially controlled compositions(SCC).sitnultaneous microstructure and Inacrostructure control.Structural ceramic componentsWe have recently Inodified the 3DP process under ARPA/ONR support to produce almnina components which are greater than 99.3% dense after fuing and have average flexural strength of 360 MPa[7].

Sintered silicon nittide parts with average flexlu'al strengths of 570 MPa have also been Inade by 3DP[8]. The basic elelnents of the modified process are to spread sublnicron almnina powder and plint latex binder. Green pm'ts m'e isostatically pressed and sintered to densify the cOlnponent. The polylneric binder is removed by thennal decOlnposition plioI' to sinteting.For example, gradual blending of Inatelials with different coefficient of thennal expansion (CTE) is possible with 3DP. Metal parts with sUlfaces that are cermnic-rich can be Inade Internal residual for high telnperature applications or for ilnproved wear resistance. stresses can be Inodified with second phases to strengthen or toughen Inaterials. These m'e just few of the possible applications of 3DP in SCC.SCC for structural matelials can have many benefits and unusual features.The CUl1'ent 3DP process involves printing of continuous binder jet onto powder bed.

Consequently, the ability to prepare a stable dispersion becOlnes a detennining factor for the size and mnount of pmticulate second phase in the final part. Agglolneration in the binder systeln causes improper nozzle perfonnance and reduces ctinlensional acclu'acy and182resolution. Khanuja has established a preparation and filtration schelue for fine alumina dispersion. [9] In his study, electrostatically stabilized submicron alumina dispersions were tested. Alumina slun)' with up to 40% in volume has been successfully dispensed through typical nozzle for 3DP. Calculations indicate that the maxilumu volmue fraction of second phase that can have COlue through nozzle with the cunent 3DP process is 30 vol% of fInal patts.

Calculations indicate that the maxilumu volmue fraction of second phase that can have COlue through nozzle with the cunent 3DP process is 30 vol% of fInal patts. However, modifications to the process routes such as the luultiple plinting cycle per layer and use of porous powder bed luay allow the fabrication of pat1S with lat'ger amounts of second phase.Zirconia toughened alumina (ZTA) system was chosen as the first luodel composite systelu and this system requires only 15 to 20% zirconia dispersed in almmna to attain optimum sttength[lO]. Fine zirconia particle size and nanow size distlibution is also These requirements match precisely with the capabilities of current 3DP required. technology. Simple calculations suggest that one only needs to prepat'e a zirconia sImi)' of 12.5 vol% solids loading to fabricate parts with 15 vol% zirconia.

The fracture energy required to propagate the crack front increases due to the n'ansfonuation. Vohuue expansion of sn'engthening[ll]. Expanded 1110noclinic precipitates cause cOlupressive stress build up on the crack tip. Surface glinding of a ZTA part induces the phase ttansfonuation near the sluface and leads to residual cOlupressive techniques developed to fabricate ZTA with thenually inelt sn'ess[12]. Vat'ious treatluents in destabilizing atluosphere and sm'face compressive sUlfaces include heat coating steps[13-18]. Often these processes require prolonged heat treatluents and present manufactlu'ability concerns. 3DP technology can be ilupleluented to fabricate ZTA with built in compressive layer by exercising the spatial conn'ol of cOluposition.thetetragonalthe precipitates also causes3DP can be used to deposit zirconia sluny onto a bed of ahumna powder to create a green ZTA part with conn'olled zirconia distlibution.

Realization of this concept is illustrated in Figure 1. Cross sections of each line plinted with 5 vol% zirconia demonstrates 3DP's ability to control level. Although Figure 1 shows the first ZTA part with only one zirconia cOluposition, two or more nozzles can be used to deposit zirconia slunies of different stabilizer contents. For example, pure zirconia slun)' can be plinted in the regions where spontaneous n'ansfonuation to luonoclinic phase is desired. Yttria doped zirconia luay be deposited in the rest of the part where the presence of luetastable tetragonal phase is prefened. Thus, ZTA parts with user specified residual stress profile can be fabricated by luanipulation ofthe zirconia contentin luicrostructural183the 3DP process. such as those with sUlface compression stresses or with a duplex ZTA microsturcture.Figure 1:Schematic of the first 3DP-derived ZTA part is shown on left. 5 vol. % Zr02 sluny was printed on Ah03 powder.

5 vol. % Zr02 sluny was printed on Ah03 powder. A cross-section of the CIPed and fired sample is also shown. Each bright feature is a cross-section of a printed line.Advanced nlaterials for medical devicesCOlnputer-derived microstluctures such as for SCC and functionally gradient Inaterials (FGM) have several medical applications such as drug delively devices and structures for tissue regeneration. Our focus has been to delnonstrate use of biOlnedical polylners with the 3DP process for both of these applications[6,19]. Attempts to control drug release rates of implantable and oral drug delively devices (DDD) by Inanipulating the geometric structure of the devices have been reported. Cylindrical rods with intelual pie shapes, inwardly-releasing hemispheres, and cylindrical donuts, represent attempts to affect release rates by modifying macroscopic shape[20]. These cOlnplex shapes are, however, not easily produced by conventional methods.

These cOlnplex shapes are, however, not easily produced by conventional methods. Conventional manufacturing techniques also tend to produce homogeneous blends of the drug and Inattix material[21]. The ability to control local composition, microstructure, and spatial distribution of drugs, makes 3DP a viable processing technique for construction of DDD. Prescriptive dosage regitne DDD with controlled gradients in drug composition and Inatrix microstructure can be created by 3DP. The mechanislTI of resorption can be controlled by selecting the appropriate binder material.Tissue regeneration devices are structures used as the matrix for tissue growth during healing. These devices are proposed for use in cases where natural tissues have been damaged by disease or traUlTIa. Generally these Inaterials are resorbable and ultitnately become natural tissue, such as skin, cm1ilage, bone, and organ tissue.

Generally these Inaterials are resorbable and ultitnately become natural tissue, such as skin, cm1ilage, bone, and organ tissue. Medical procedures based on this approach are being developed for a number of applications. NUlnerous authors[22-25] have reported that optilnal devices require control of their structure over several levels of structural hierm'chy, including gross macroscopic shape, oriented pores184and channels, and microporosity. The ability to preferentially prOlnote cellinigration and angiogenesis, for example, can be accOlnplished by directing nutrient delivery in a complex cell seeding structure. Precise structural control beyond randOln lnicroporosity has not been achieved by conventional processing Inethods. The potential to also intiInately control the orientation and placement of porous channels and the overall Inacroscopic shape of a device makes 3DP an ideal process for producing tissue regeneration devices.

Below we describe several devices that delTIOnstrate reticulated structures.Polyethylene oxide (PEO) and polycaprolactone (PCL) are selected as Inatrix Inaterials for trial experiments and methylene blue and alizarin yellow were used as drug models. The dye release kinetics is controlled by either specifying the position of the dye within the device or by controlling the local Inicrostructure with the 3DP process. The dye release rates of the devices are characterized by placing single devices into individual beakers filled with 10 ml water at rOOln temperature (20°C). Sample solutions are collected, and fresh 10 ml of water is replaced at appropriate tilne intervals. Absorbance for Inethylene blue (664 mn) and alizarin yellow (353 mn) is Ineasured for all saInples on a DU-640 spectrophotometer. Spectrophotolnetric analysis of dye release yielded reproducible results.

Spectrophotolnetric analysis of dye release yielded reproducible results. Three examples are described below.The Inodel device used for this study is designed to delnonstrate regulation of drug release profiles by controlling position, cOlnposition, and Inicrostructure. A diagraIn of the device is shown in Figure 2a. The top and bottOln of the tabular device are cOlnposed of relatively nonresorbable PCL while the interior layers are cOlnposed of PEO bound by plinting binders solutions so as to fonn perpendicular walls in the interior of the device. Dyes are deposited in selected locations within the device during the build procedure. The top and bottoln sheets of dense PCL serve as baITiers against dye diffusion since hydrolytic degradation of PCL occurs Inuch slower than that of PEa. Therefore dye diffusion during resorption is confined to the plane of the device since the top and bottOln aI°e only slowly resorbable.

Therefore dye diffusion during resorption is confined to the plane of the device since the top and bottOln aI°e only slowly resorbable. Figure 2b shows the diagraIn and release profile of a device with symmetrical spatial distribution of the two dyes and unifonn matlix gradients. All walls in this device are identical since they are constructed with the san1e binder cOlnposition, under the SaIne printing conditions. As expected, the blue and yellow dyes are released at identical rates to within the statistical error of the experilnent.One of the plinting paratneters which can be tnodified to regulate release rates is the inter-line spacing of the lines which make up the walls. The walls of the device in Figure 3a are chemically identical, but physically different. Six solid walls are built in one direction by printing four lines closely together. Then six walls are constructed in the perpendicular direction by printing only two lines, separated frOln each other by unprinted powder.

Then six walls are constructed in the perpendicular direction by printing only two lines, separated frOln each other by unprinted powder. All walls are constructed to the SaIne total thickness. The device shown in Figure 3a demonstrates that for walls of equal thickness and cOlnposition, Inodification of internal wall tnicrostructure can affect the release rate. Here, the dyes lilnited by solid walls (4 connected lines per wall) are released at a slower rate than dyes lilnited by selni solid walls (2 lines per wall, spaced apart by unprinted PEa powder). It should be noted that other printing paraIneters, such as printhead traverse velocity, binder flow rate, and 3D Ptinting also feature layer thickness, can also be contl'olled to Inodify release rates.

Release profile is affected by controlling spatial distribution of dyes.186, ...,., VeliON---.!>- Blue101210121414Figure 4:Figure 5:in chloroform; Binder per Line Length: x 10-4 cn13/cnl; Powder: NaCI 75-150 Uln; Layer thickness: 150 urn; These devices were leached in water for 24 hours after printing and drying to produce the n1icroporosity visible.Binder: 30 wt~) acid I110difiedBinder: 26.5 wt% acid n10dified PCL in chloroform; Binder per Line Length: 3.33 x 10 -4 cn13/cm; Powder: 250 urn; This device was printed horizontally (vertical direction in figure) to produce regions with different density. The figure shows a cross section of this device. Note the outer channels that are not completely open due to bleeding of the binder into those regions.75-150 urn; Layer thickness:'187shown in Figure 3b, the walls are identical in both directions.

The difference in release patterns for the two dyes can only be attributed to the difference in spatial disnibution of the dyes within the device. This release profile exhibits two Inaxilna ill the Inethylene blue release rate and a relatively constant release rate of alizalin yellow.Interconnected pores on the order of 50-250 !-un within a device provide a structure for cell proliferation, migration, excretion of Inauix Inatelials, and blood vessle fOlmation. Pores with width greater than approxilnately 100 Jlm can be built directly with the 3DP process. Building a device with smaller pores requires the inclusion of a leachable component. This leachable component can then be relnoved after plinting. An exalnple of the microporosity created by this process is shown ill Figure 4. These pores are of the correct size for viable cell function and angiogenisis. The powder bed matelial, ill this case NaCI, has been cOlnpletely leached out of the final device.

These pores are of the correct size for viable cell function and angiogenisis. The powder bed matelial, ill this case NaCI, has been cOlnpletely leached out of the final device. Thus, the device is the fITst 3DP device of any kind to consist only of material delivered through the printhead. Other 3DP devices incorporating natural Inatelials, such as bovine bone, have been constructed. In this case, bone powder was mixed with PCL powder. A solvent polYlner solution was used as the binder.A cross section of a macroscopic device constructed with PCL and polYlner solution is shown in Figure 5. All of the intelior channels are open, but SOlne of the channels near the dense, high binder delivery regions are obstructed. This is likely due to the bleeding of some of the binder into the void regions.Final perspectivethe unusual This blief report capabilities of solid-free fonn fabrication Inethods.

This is likely due to the bleeding of some of the binder into the void regions.Final perspectivethe unusual This blief report capabilities of solid-free fonn fabrication Inethods. We have already delnonstrated unprecedented conn'ol of microstructure and local cOlnposition using the 3DP Inethod, but other SFF technologies will undoubtedly be used along silnilar lines. Never before have matelials and product designers had the ability to specify structure on the Inacro and micro level. The nlunber of possible applications seems innmnerable. Indeed, our Inost difficult task may be to select applications to develop first. An ilnportant consideration for selection of the appropriate process will, however, be the production rate. COlnponents with unique microstructures will only be useful if thousands of parts can be Inade. 3DP may be uniquely qualified to address this .lnanufactuling issue.

COlnponents with unique microstructures will only be useful if thousands of parts can be Inade. 3DP may be uniquely qualified to address this .lnanufactuling issue. Multijet plinting is now possible on 3DP machines as has been den10nstrated at Soligen£26l and at MIT. This has dramatically increased the production rate of the 3DP process. C0l11lnerciai ink-jet print heads are available with thousands of individually conu'olled jets. Thus, future 3DP for prototyping. machines may be closerrepresents SOlne of the first experitnents to exploitto production toolsratherthan tools188AcknowledgmentsThis work was supported by ARPA/ONR, Johnson & Johnson Developluent Corp. and Therics Inc. B. Wu acknowledges support frolu an Institutional Dentist Scientist Award by N.LH. and Harvard School of Dental Medicine. The authors express their gratitude to J. Serdy and the M.LT. 3D Plinting group for technical assistance. Finally, the authors wish to thank Dr. Walter Flatuenbaum and Mr.

The authors express their gratitude to J. Serdy and the M.LT. 3D Plinting group for technical assistance. Finally, the authors wish to thank Dr. Walter Flatuenbaum and Mr. Tucker Swan for luany interesting technical discussions.References1.E.M. Sachs, M.J. Cima, P. Williatus, D. Brancazio, J. Conlie, "Three-Diluensional Printing: Rapid Tooling and Prototypes Directly from a CAD Model"; J. Eng. Ind., 114 pp. 481-88 (1992).2.E. Sachs, M.J. Chua, J. Bredt, A. Curodeau, "CAD-Casting: The Direct Fabrication of Ceratmc Shells and Cores by Three Diluensional Plinting," Man. Rev. 5 [2] pp. 117-126, (1992).3.M.J. Cima and E.M. Sachs, "Three Dituensional Printing: Fonu, Materials, and Perfolmance," in Proceedings of the Solid Freeform Fablication Syluposium Edited by J.J. Beatnan, H.L. Mat"cus, D.L. (8/12/91-8/14/91, Austin, TX). Bourell, J.W. Barlow, and T. Crawford. University of Texas, Austin, TX pp. 187 94.4."Metal Parts Generation by Three S. Michaels, E.M. Sachs, M.J.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperConceptual Design for Additive Manufacturing: Lessons Learned from an Undergraduate CourseScott M. ThompsonDept. of Mechanical & Nuclear Engineering, Kansas State University, Manhattan, KS 66506Email: smthompson@ksu.eduAbstractDesign for additive manufacturing (DfAM) guidelines continue to emerge and evolve as various additive manufacturing (AM) technologies, and the knowledge of their associated end- users/designers, matures. This work summarize important pedagogical and technical lessons learned from the conceptual re-design of several, diverse parts/assemblies submitted by ~50 undergraduate students cognizant of recent DfAM strategies and guidelines. All students were enrolled in a traditional, semester-long metals AM course designed by the author herein.

All students were enrolled in a traditional, semester-long metals AM course designed by the author herein. Students were instructed to select an existing, metallic product and provide a conceptual redesign of that product for subsequent, effective laser-powder bed fusion (L-PBF). Students were instructed that the redesigned concept should have enhanced functionality/specifications and consist of features (e.g. thin walls, bore diameters, etc.) that can be fabricated with minimal risk via current L-PBF systems. The presented results include the types of parts that attract an ‘AM redesign’ effort and the most popular AM-enabled detailed design decisions made. To encourage a more detail- inspired design, DfAM from post-manufacturing topics were presented ‘backwards’; considerations to conceptual design while considering emerging design rules and heuristics.

Results indicate that students can become preoccupied with DfAM rules to a point where the design failure modes are not properly accounted for.KEYWORDS: Design for additive manufacturing, education, training, conceptual designIntroductionAdditive manufacturing (AM), and especially laser-powder bed fusion (L-PBF), has broadened the design space for engineers leading to the production of more material-efficient products for various applications. However, the disruptiveness of AM has led to a disconnect between engineers, both junior and senior, and the effective means for addressing when AM should be employed for a product and the best approach to design for AM (DfAM).The increased design space offered by AM challenges engineers to think differently. They must consider out-of-norm concepts and shapes that were once infeasible to manufacture. There has been research into how to inspire or guide engineers to take advantage of AM via new designs.

They must consider out-of-norm concepts and shapes that were once infeasible to manufacture. There has been research into how to inspire or guide engineers to take advantage of AM via new designs. For instance, Rosen [1] has provided some methods for engineers to adopt for considering new designs via biomimicry and cellular structures. Booth et al. have provided visual DfAM worksheets to assist students and technicians in identifying AM constraints and important DfAM rules [2].56Engineers involved with design/manufacturing must also remain ‘grounded’, and cognizant of the AM process and its limitations. The process-structure-property-performance relationships inherent to the designed/actual product and its manufacturing method must be considered. There have been studies into the pedagogy of DfAM and its constraints [1–11].

There have been studies into the pedagogy of DfAM and its constraints [1–11]. Many of these works have aimed to formalize DfAM, provide DfAM heuristics, determine new training methods for undergraduates and practicing engineers, determine effects of part orientation/geometry on tolerance and surface roughness, and more. The main challenge with DfAM is that the AM technology itself is evolving rapidly. It is possible for some DfAM heuristics to become dated quicker than the publication process inherent to DfAM textbooks and papers.BackgroundAn undergraduate/graduate course called ‘Additive Manufacturing of Metals’ was taught during the spring semester of 2019 at Auburn University. The course was a survey-type, traditional lecture-style course. Although the course allowed for graduate students to enroll, the majority of students were mechanical engineering undergraduates. Total enrollment was 48 students (including 3 graduate students).No textbook was adopted for the course.

Total enrollment was 48 students (including 3 graduate students).No textbook was adopted for the course. Instead, students were exposed to the methods and findings from various academic papers Refs. [1–11], online videos, guest speakers, and the experience of the author. A particularly useful paper by Pradel et al. was employed for discovering and teaching DfAM papers based on their detailed literature survey [3].The first half of the semester was devoted to the major AM types, case studies and terminology. Throughout the semester, students were provided with “In the News” updates before lecture that summarized key advancements in the field and industrial use of AM. The second half of the semester focused on DfAM – including detail, embodiment and conceptual design heuristics and methods as they pertain to AM – especially metals AM.

The second half of the semester focused on DfAM – including detail, embodiment and conceptual design heuristics and methods as they pertain to AM – especially metals AM. Some other topics discussed with DfAM included: a Review on basic design and optimization principles/methods; Post- manufacturing techniques (wire EDM, heat treatment, surface finishing, etc.); AM process chain, thread for AM; Computer-aided design considerations for AM; and material digital selection/substitution. Students were taught DfAM by using a ‘backwards’ approach used by Pradel et al. Meaning, students were first introduced to concepts & rules/heuristics related to post- AM operations, followed by (in order): part programming, detail design, embodiment design, conceptual design, and requirements. By teaching detailed design first, students were being steered more toward embodiment design approaches which depend more on material, details, assembly and functionality [12].

By teaching detailed design first, students were being steered more toward embodiment design approaches which depend more on material, details, assembly and functionality [12]. This presentation method is believed to make students more likely to consider post-AM and programming/CAD issues during initial design – as these final two steps can influence DfAM more so than other design-to-manufacturing efforts.For the final month of the course, students were assigned a project that challenged them to take any product and re-design it for AM as to enhance the product. All students were instructed to consider several rules for DfAM and to re-design while being cognizant of these rules. They were to identify features that hinder the additive manufacturability of the legacy part they considered. Students were also tasked in determining possible failure modes of the AMed part. Sketches of the new AM product were asked to be provided.

Students were also tasked in determining possible failure modes of the AMed part. Sketches of the new AM product were asked to be provided. Students had the option to work with a partner on the design project. The major DfAM rules students considered included: smallest feature, functional design, surface quality, dimensional accuracy (as-printed), material availability,57time, powder entrapment and de-powdering, size constraints, mechanical performance, part orientation in AM chamber, support structures, and more.ResultsSeveral solutions were submitted as part of the course project, however, only a select few are presented to demonstrate the conceptual DfAM approach employed by undergraduate students. It is also the goal to demonstrate how students addressed the limitations/opportunities of metals AM.

It is also the goal to demonstrate how students addressed the limitations/opportunities of metals AM. It is worth noting that some of the designs not provided here include a vegetable peeler, carburetor main body, metal baseball bat, golf club, hammer head, axle shafts, piping elbows, iron soleplate, and a dock cleat. The effort dedicated by the students on the project assignment varied. Students were generally graded on whether instructions were followed and for completeness.Cattle brandA group of two students identified a cattle brand as a candidate for re-design for AM. The design pitch indicated that there is an opportunity to make unique/custom brands with features less harmful to livestock. As shown in Fig. 1, the students called out regions along a traditional cattle brand that require careful attention for DfAM.Fig. 1 Legacy cattle brand design with regions challenging the effective AM labeled.Label descriptions for Fig. 1:1.

1 Legacy cattle brand design with regions challenging the effective AM labeled.Label descriptions for Fig. 1:1. Lack of fillets in design lead to stress concentrations and resolution difficulties. 2. Even though the main loading will occur parallel to the build (Z+) direction, any accidental loading in the XY plane, perpendicular to build direction, will produce excessive transverse shear in these thin beams.3. With a Z+ build direction, most of this brand will require excessive amounts of support material. Which will be impossible to access/remove during the post processing steps. 4. Voids in the design will behave as sites for fatigue crack nucleation and propagation. 5. Overhangs in the part will require support.586.In general this brand is very tall for AM processing, this is a major limiting factor to its additive manufacturability because there is a strong positive correlation between cost and Z axis height.Based on the ‘problem areas’ shown in Fig.

1, a new cattle brand design, with enhanced features, was proposed for AM as provided in Fig. 2. Note that the letters on the AM cattle brand were arbitrarily chosen, but represent a sequence of letters/numbers that may be used for branding.Fig. 2 Proposed cattle brand design for AM.ZipperOne student challenged the traditional zipper design – especially with respect to its extended tab feature which can be difficult to manage for some people. Of course, such a device would be infeasible to fabricate with AM economically, yet the provided solution elucidates more on the DfAM approach. Sketches of the re-imagined zipper for AM are provided in Fig. 3.l'T. , f , ' A 1 1 I l l . , ~ : J c V:e .,(f ), J L, "•♦ A l l , r .< ' , . . . , I f . , . J( j Ji J - r - v [ _ y/ H "/ 1 . ,q J , J ~Fig. 3. Conceptual design sketches for re-design of a zipper for metals AM.59The student indicated that the primary reason for re-designing the zipper was usability.

,q J , J ~Fig. 3. Conceptual design sketches for re-design of a zipper for metals AM.59The student indicated that the primary reason for re-designing the zipper was usability. A noted challenge would be redesigning zippers to not have any hard 90° transitions between faces. In addition, the printing a single part that has interlocking pieces is a challenge. As discussed by the student the first alternate design (in Fig. 3) can be arranged in such a way, that if oriented with y- dimension perpendicular to the building plane (so that x and z are within the 2D build plane), the tab can be supported free of the zipper, allowing for easy support removal. The student mentioned that the AM zippers were all designed to have their y-dimension perpendicular to the build plane for the ease of support removal.

The student mentioned that the AM zippers were all designed to have their y-dimension perpendicular to the build plane for the ease of support removal. The student was aware of surface roughness effects and discussed how the orientation of the zippers on the substrate could help control this feature.LEGO® diesLEGO dies and other polymeric dies are not currently produced via AM because of the low tolerance and demand for a pristine surface. As shown in Fig. 4, the student addressed some major challenges in using metal AM for fabricating LEGO dies. The student was clearly aware of application environment and how this would impact the material selection and design. As mentioned by the student, thermal cycle fatigue resistance is an important property of the LEGO die as it would experience ~200 ºC thermal cycle for each use.

As mentioned by the student, thermal cycle fatigue resistance is an important property of the LEGO die as it would experience ~200 ºC thermal cycle for each use. To survive this use a metal with high thermal conductivity will be more beneficial so that the heat can be transferred evenly through the part and into the surroundings. The student mentions stainless steel as meeting this criterion better than other common AM metals. The surface roughness, minimum feature size and corner radii of the die were mentioned as major challenges for ensuring a successful AM LEGO die.Surface AnishFig. 4. A photo of a LEGO die with challenges for AM pointed out.Bicycle stemA team of two students worked on the re-design of a typical bicycle stem for AM. A bicycle stem is a bracket that connects the handlebars to the headset on the frame. A common bicycle stem with noted challenge areas for DfAM are provided in Fig. 5. Solutions to the problem areas marked in Fig.

A common bicycle stem with noted challenge areas for DfAM are provided in Fig. 5. Solutions to the problem areas marked in Fig. 5, as provided by the students, is also provided below Fig. 5. This group was effective in considering the actual build plan for the AM of the bicycle stem, and the envisioned orientation/placement of the part on a L-PBF build plate is provided in Fig. 6.60Difficult support remova l from middleTh reads are tooPart volume is not optim izedsmall/have knifeSmoot h surface finish requiredPart is made of alum inum35 Degree OverhangLarge Diameter Hole acts as overhang when printed on sideFig. 5. A photo of a bicycle stem with challenges for AM pointed out.Some notes from the students with regard to the challenges in fabricated the bicycle stem via AM are provided below.1. This part will need to be printed on the side to avoid having the large 35-degree overhang.

This part will need to be printed on the side to avoid having the large 35-degree overhang. This will also act to support the load along the print plane to prevent delamination.2. The frame-side hole will require support between the diameter. 3. The part will use compatible metals such as titanium and steel. It is necessary to reconsider if the material can withstand the environment, weight, and strength.4. Because threaded holes are not manufacturable via AM, post-AM tapping will be required to create the threads.5. To fix both the optimization issues and the support removal issue, a topology optimization method and FEA method could be employed to use less material in the shaft. This would reduce the weight of the part and allow holes for support to be removed easier.6. The inside of the holes will need post-AM machining or sanding so that the handlebars and headset don’t get scratched61Fig. 6.

The inside of the holes will need post-AM machining or sanding so that the handlebars and headset don’t get scratched61Fig. 6. A sketch of how the bicycle stem would be positioned within a L-PBF chamber for fabrication. Some notes are provided by students in the sketch.Fish hookOne student provided a conceptual design for a new fish hook that employs more curvature and a j-type top portion to allow for one to wrap a fishing line around the hook rather than using a hole. This concept is shown in Fig. 7. The provided a simple new functionality of a fish hook based on considering the advantages of AM, yet was still considerate of the limitations as provided by the use of support structure.r;,1,t ~ill. ~ br•«.______. ,! IA"' k.]JFig. 7.

~ br•«.______. ,! IA"' k.]JFig. 7. Student sketches demonstrating (left) a typical fish hook, (middle) an AM fish hook with support brace concept, (right) a hole-less fish hook concept for AM.Aerospace rotorOne student investigated the challenges in adopting AM for the fabrication of an aerospace rotor (Turbocam). This specific rotor is one of the first components inside a gas turbine, specifically the compressor. The student noted how the rotor is a critical component to gas engine operation and public safety and that common failure modes are erosion and fatigue failure due to high cyclic loading operations. As shown in Fig. 8, some common AM design rules are addressed at pointed- out locations on the rotor. The student was aware of the build envelope constraint for the selected rotor; however, the student noted alternate AM methods (to L-PBF) such as Big Area AM (BAAM). A re-design of the rotor for AM, as provided by the student, is shown in Fig.

A re-design of the rotor for AM, as provided by the student, is shown in Fig. 9.62Rules Violated Maximum Prut SizeEx lruiation Th.is rotor feanires ru1 outside diameter of 23 inches. Most metal additive mrurnfac.tt11·in2 machines do not allo,; for th.is big of a build volmne.Annotated PictureMininnun Wall th.ickness Th.is rotor featt11"es a total of26 blades each o fve1y small thickness.Minimum Feature sizeThis rotor features ve1y small chrumels for the blades. These chaimels ru·e used to increase the outlet pressure from the rotor.Knife EdgesAt the bottom side of each blade there ru·e shaip edges.Faces Requiring SupportAt the end of the pa1t there ru·e retaining feattu·es that have overhanging feanires.Fig. 8. AM challenges associated with an aerospace rotor.Fig. 9. Proposed re-design of an aerospace rotor for AM.63All drawings and reports were compared against each other in order to determine general DfAM approaches adopted by the students.

9. Proposed re-design of an aerospace rotor for AM.63All drawings and reports were compared against each other in order to determine general DfAM approaches adopted by the students. It was clear that students set higher priority to ensuring their re-designed parts were additively-manufacturable. Re-designed parts did not necessarily have better features for enhancing product lifetime (failure resistance). This could be attributed to the assignment time constraint hindering any ability for students to predict and compare part performance during application. Students re-designed parts for functional improvement while not altering design concepts drastically. By teaching detail-driven DfAM, students appear to have focused more on embodiment of their legacy products rather than radical new concepts. In their focus toward embodiment and functionality, the vast majority of selected parts for re-design were not assemblies in need of consolidation, rather, they were single-bodied parts.

In their focus toward embodiment and functionality, the vast majority of selected parts for re-design were not assemblies in need of consolidation, rather, they were single-bodied parts. Students were clearly able to use common DfAM rules and to identify potential geometric alterations in legacy parts for improving additive manufacturability.Conclusions & RemarksBy inspecting the sketches and write-up of senior-level mechanical engineering undergraduate students that focused on the the conceptual, embodiment and detailed designs for the AM of legacy, everyday parts, it was possible to take a ‘snap shot’ of how students utilize DfAM heuristics and rules. All students were presented with several DfAM concepts/approaches over half a semester. The below are some major conclusions and remarks from this study.1. Teaching the DfAM stages backwards encourages students to focus more on embodiment rather than concepts.

The below are some major conclusions and remarks from this study.1. Teaching the DfAM stages backwards encourages students to focus more on embodiment rather than concepts. This could be beneficial for design efforts focused more on product enhancement.2. Although part consolidation is an advantage provided through AM, the vast majority of students did not select legacy products with multiple joints/parts.3. Students can easily apply DfAM rules for aiding the re-design of parts after they are provided with lists and worksheets from the open literature. Students can become preoccupied with these rules and provide less attention to designing against product failure during application. 4. The assignment, although limited in scope and time, provides a valuable opportunity for students to rehearse important DfAM rules and to critically evaluate the functionality of current products.

4. The assignment, although limited in scope and time, provides a valuable opportunity for students to rehearse important DfAM rules and to critically evaluate the functionality of current products. It is recommended that DfAM rules involve multiple levels of specificity – with their being broad, agnostic sets of rules that can apply to a particular AM method, and then more detailed rules for specific AM systems and materials. More agnostic DfAM rules are highly needed to combat their antiquation as AM technology continues to rapidly evolve. The use of ‘cheat sheets’/worksheets and lists with visual graphics should aid the DfAM rule retention and usability moving forward.AcknowledgementsA portion of this work was conducted while the author was a faculty member at Auburn University.

All students were made aware that their submitted work could be included in a research article, and only the sketches/photos generated by consenting students have been included in this paper.64References[1]Rosen, D.W., 2007. Design for additive manufacturing: A method to explore unexplored regions of the design space. Proceedings of the 18th Solid Freeform Fabrication Symposium (SFF 2007), 6–8 August 2007, Austin, TX, USA, 402–415.[2]Booth, J. W., Alperovich, J., Chawla, P., Ma, J., Reid, T. N., and Ramani, K., 2017, The design for additive manufacturing worksheet, ASME. Journal of Mechanical Design, Vol. 139, No. 10, p. 100904.[3]Pradel, P., Zhu, Z., Bibb, R., Moultrie, J., 2018, A framework for mapping design for additive manufacturing knowledge for industrial and product design, Journal of Engineering Design, Vol. 29, No. 6, pp.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperConceptual design for assembly in the context of additive manufacturingS. Yang* and Y. F. Zhao*Department of Mechanical Engineering, McGill University Montreal, Quebec, Canada H3A 0C3AbstractAs additive manufacturing (AM) emerges as an end-of-use product manufacturing process, design for additive manufacturing (DFAM) as a new design philosophy receives more and more attention. However, current DFAM research focuses on downstream and part-level design activities such as structural optimization and design rules. Design freedom enabled by AM such as, part consolidation and function integration has not been fully investigated.

Design freedom enabled by AM such as, part consolidation and function integration has not been fully investigated. These design freedom forces designers to rethink about assembly-level design so as to embrace integrated functionality. To understand how to integrate AM characteristics into design process, three questions are investigated: 1) why does conceptual design need to be redone for assembly? 2) what has changed by AM in design concept generation? 3) how to do conceptual design in AM context? Afterwards, a conceptual design framework is proposed to aid design flow management. In the end, a throttle pedal redesign case is demonstrated as verification of the proposed design framework. Keywords: assembly-level DFAM, conceptual design, additive manufacturing, functional design, FBS modeling1.

Keywords: assembly-level DFAM, conceptual design, additive manufacturing, functional design, FBS modeling1. IntroductionAdditive manufacturing (AM) is defined as “a process of joining materials to make object from 3D model data, usually layer upon layer, as opposed to subtractive manufacturing methodologies" [1]. This layer-wise manufacturing process makes it possible to manipulate material distribution and composition without any other tools. The design freedom enabled by such a technique includes shape complexity, hierarchical complexity, material complexity, and function complexity [2, 3]. All these new features of AM force designers to incorporate AM-enabled design freedom into the design stage and rethink about how to make a better engineering design. However, as Yang et al. [3] and Laverne et al.

However, as Yang et al. [3] and Laverne et al. [4] pointed out, current Design for Additive Manufacturing (DFAM) research focuses on part-level and downstream design activities such as structural optimization and design rules without the possibility to be extended for assembly-level applications. These applications mainly refer to part consolidation and function integration, which helps to reduce part count and product weight and size, to eliminate assembly difficulty, and to improve design robustness. All these benefits make assembly-level DFAM receive more and more attention.In comparison with part-level DFAM, assembly-level applications have to deal with relations between components in order to decide if two components can be consolidated or not. Moreover, part consolidation will lead to function integration and even a coupled design which is not favoured by Axiomatic Design [5]. How to make the trade-off in design complexity is unclear.

Moreover, part consolidation will lead to function integration and even a coupled design which is not favoured by Axiomatic Design [5]. How to make the trade-off in design complexity is unclear. Furthermore, assembly-level applications require more complete information than the part-level counterpart. For example, the boundary of the system of interest should include necessary void information within the convex hull of the chosen components. Apparently, all these three aspects1932are intensively related to the conceptual design stage of the system of interest. To upgrade DFAM philosophy into the assembly-level design to embrace the advance in function integration and part consolidation, a comprehensive understanding of how a product is developed is necessary, especially the conceptual design stage. Besides, conceptual design has its practical value because the initial stage of design accounts for 80%-90% of the product success [6].

Besides, conceptual design has its practical value because the initial stage of design accounts for 80%-90% of the product success [6]. Unfortunately, existing Design Theory and Methodology (DTM) cannot support the need of taking advantages of AM- enabled design freedoms [3, 7]. From conventional design philosophy point of view, it is advocated to design simply and dissect products to ease assembly operation and reduce manufacturing difficulties.

From conventional design philosophy point of view, it is advocated to design simply and dissect products to ease assembly operation and reduce manufacturing difficulties. Therefore, the problem of conceptual design for assembly in the context of AM cannot be solved by conventional DTM.In order to shed light into conceptual design and provide a methodological way to applyAM-enabled design freedom, three questions are to be investigated in this paper.Question 1: how does existing conceptual design philosophy work? • Question 2: what are the challenges in the context of AM? • Question 3: how to do conceptual design with appropriate AM knowledge input?These questions are discussed in Section 2, 3, and 4 respectively. A function-behavior-functional entity-functional feature (FBFF) design synthesis method is proposed to answer Question 3. Afterwards, an example of a throttle pedal is investigated in terms of conceptual redesign.

Afterwards, an example of a throttle pedal is investigated in terms of conceptual redesign. Finally, the paper is wrapped up with conclusions and further research.2. Existing conceptual design philosophyIdeally, product design process needs to understand the nature of decision making and the constraints in different design phrases. In this section, several design methodologies involving the concept of conceptual design are discussed. These methods include General Design Theory (GDT) [8], Systematic Design approach [9], VDI 2221 [10], and Axiomatic Design [5]. In this paper, function is defined as a description of behavior abstracted by human through recognition of the behavior in order to utilize the behavior [11].2.1 General Design TheoryGDT deals with concepts that only exist in our recognition and it uses set theory to explain knowledge manipulation in the design process. It has a basic manifesto that knowledge can be mathematically formalized and operated.

It has a basic manifesto that knowledge can be mathematically formalized and operated. Therefore, design process is a mapping from function space to attribute space. As shown in Figure 1, function space is a complete set of function entities. Therefore, design solutions can be regarded as the results of logical operations of entities concepts. Afterwards, an entity that can fulfill these design requirements are enriched with abstract concepts. If there is no design solutions found, then knowledge in the first step is incomplete and it becomes the core step of design synthesis. Otherwise, if such an entity concept as the design solution is found, the solution is mapped from function space to attribute space. Then, the neighborhood of the entity in the attribute space is analyzed to obtain attributive information for production. These attributive information includes shape, geometry, material, etc. GDT helps to understand the design process from an entity point of view.

These attributive information includes shape, geometry, material, etc. GDT helps to understand the design process from an entity point of view. Given ideal knowledge, design will be the results of logical operation of function entities and attribute entities.1933Figure 1 Design process in ideal knowledge [12]2.2 VDI 2221 and Systematic design approachThe most well-known design method in practice is the Systematic Design approach by Pahl and Beitz [9]. In this design method, design process is comprised of four phrases: planning and clarifying, conceptual design, embodiment design, and detail design. In the conceptual design phrase, design requirements are interpreted into the principle solution and then, the principle solution is developed into a design concept (Figure 2 a). Similarly, VDI 2221 [10] shares the same procedure but it makes the transition from design requirements to layouts more explicit as shown in Figure 2 b.

Similarly, VDI 2221 [10] shares the same procedure but it makes the transition from design requirements to layouts more explicit as shown in Figure 2 b. In VDI 2221, the counterpart of conceptual design covers steps of “determine functions and their structures”, “search for solution principles and their combinations”, and “divide into realizable modules”. Besides, Systematic Design approach models function as input-output flow of energy, material, and signal. This function model helps to do functional decomposition based on the causal relation between functions.2.3 Axiomatic designIn Axiomatic design, the best design solution is supposed to satisfy two axioms [5]: maximum independence of functional elements and minimum information content. The design process (Shown in Figure 3) is divided into four domains: customer domain, functional domain, physical domain, and process domain.

The design process (Shown in Figure 3) is divided into four domains: customer domain, functional domain, physical domain, and process domain. The counterpart of conceptual design in Axiomatic Design is the mapping from the functional domain to the physical domain, which also involves a vertical functional decomposition process within the functional domain. The main contribution of Axiomatic Design is providing a perceptual way to understand a general design process and a mathematic way to evaluate design according to the design matrix and the information axiom.2.4 DiscussionAs shown on the above, although GDT, VDI 2221, Systematic Design approach, and Axiomatic Design vary in the form of representation, they all cover two design activities in the concept developing stage: functional decomposition and design synthesis. Functional decomposition refers to process of decomposing high-level functions to lower ones until specific design solutions are available.

Functional decomposition refers to process of decomposing high-level functions to lower ones until specific design solutions are available. Design synthesis means the process of mapping functions (or functional requirements) to concrete physical forms. In practice, these two design activities are not sequential but iterative because of the imperfection of knowledge (of designers). In this sense, conceptual design can be abstracted as the enrichment of functions and the realization of functions.1934(a) Systematic Design approach(b) Recommended design process (VDI 2221)Figure 2 Systematic design approach [9] and VDI 2221 [10]Figure 3 Axiomatic design process [5] The relation between functional decomposition (or functional analysis [13]) and design synthesis is shown in Figure 4. The initial input of conceptual design is functional requirements and the ultimate output is design concepts.

The initial input of conceptual design is functional requirements and the ultimate output is design concepts. Between functional analysis and design synthesis, there may be an iterative process because functional analysis process may be aided by the hierarchy of physical product.Abstracting the conceptual design process into functional analysis and design synthesis helps to characterize the impact of AM on conceptual design. All the studied design methodologies can support developing design concepts; however, they are so general and abstract that they are insufficient to lead designers in terms of how to develop a new design concept.

When it comes to AM context, the problems will be how functional analysis and design synthesis are affected by1935AM-enabled design freedom and how to do functional analysis and design synthesis to find AM- enabled design concepts.Conceptual designFunctional requirementsFunctional analysisDesign synthesisDesign conceptsFigure 4 Main design activities in the conceptual design stage3 Challenges in the AM contextThe investigation of the impact of AM on conceptual design can be simplified as that of the influence on functional analysis and design synthesis. According to the definition of functional analysis, the main tasks will be to decompose the overall function to sub-functions until sub- functions can be realized by available physical forms.

According to the definition of functional analysis, the main tasks will be to decompose the overall function to sub-functions until sub- functions can be realized by available physical forms. Based on the classification of product innovation levels [14], most of redesign is to achieve incremental innovation rather than radical innovation because the working principles behind the new design concept is the same as the old one. Taking the design scenario of a new power system for a new car model, a design concept of solar system and a design concept of a petrol-driven system are quite different in terms of working principles; therefore, this design scenario can be regarded as radical innovation. In contrast, only changing the physical form without altering the working principles does not change the function structure.AM is more suitable to stimulate incremental innovation rather than radical innovation.

Firstly, it is difficult and unreasonable to expect companies to invest significantly to come up with a totally new design concept especially with high complexity, for example, civil aircrafts. Secondly, to pursue incremental innovation makes it easier for designers to repeat their old ways of functional design, which accelerates the product development cycle. In contrast, radical innovation is realized by a thorough change in the design concept, but AM cannot make it that far yet in most cases. For example, consolidation of horizontal stabilizer and vertical stabilizer is possible depending on the improvement in the control system instead of manufacturing techniques. Since the whole working principles are not altered, functional analysis process still follows the process of top-down decomposition and the resultant function structure will be the same as the old one.

In conclusion, AM has minor effect on functional analysis process.In design synthesis stage, the main tasks are to map low-level functions to concrete physical forms. Design synthesis is one of the most heated research topics in the design community. Design synthesis in some literatures are called functional design, functional modeling[15] or functional reasoning[16]. Since this paper emphasizes on the impact on design synthesis, it does not matter which functional modeling method is taken. Function-behavior-state (FBS) functional modeling method [16] regards function, behavior, and state as independent entities in the design process. Behavior is an aggregate of a set of states with respect to external stimuli on the basis of its1936structure and physiochemical properties. State typically refers to the multiple state of physical structure. FBS functional modeling scheme has high expressive power over the mapping from design intent (function) to physical structure.

State typically refers to the multiple state of physical structure. FBS functional modeling scheme has high expressive power over the mapping from design intent (function) to physical structure. As shown in Figure 5, a given Function A is mapped to a desired Behavior B according to human recognition abstraction. Then, Behavior B is realized by the transition of State C to D via physical structure. Thus, the mapping from a function to its corresponding physical form is accomplished. For example, given a function “generate light”, the desired behavior is “light on via electricity”, then the state will be “switch’s state is changed from off to on”. Revisiting the characteristics of AM, we know that AM is able to manipulate material composition and distribution. It means that AM is able to change morphology, topology, and material properties. In conventional design philosophy, design is limited by designers’ knowledge of existing shapes, materials, and manufacturing processes.

In conventional design philosophy, design is limited by designers’ knowledge of existing shapes, materials, and manufacturing processes. Therefore, the resultant change by AM will be the physical phenomena, that is, the mapping between behaviors to physical structures. For example, to realize the function “grab things”, the behavior will be “supporting force is larger than gravity”, then the AM-enabled feature will be “a curved surface with textures”.In conclusion, AM-enabled design freedom only enriches the mapping relations between the desired behavior and the corresponding physical structures.

The processes before this mapping including functional analysis and the recognition of desired behavior are not changed by AM.Set of functionsSet of behaviorSet of statesRecognition abstractionFigure 5 FBS modeling scheme4 Proposed design synthesis strategyAlthough it becomes clear that AM only imposes influence on the mapping from behavior to the physical form, how to stimulate designers’ creativity to come up with novel designs is still challenging. Traditionally, FBS design synthesis is realized by functional components or standard device. For example, to realize a function “store water”, corresponding physical structure will be “a container”. After that, the cross-section shape and material is chosen from the limited regular shapes and materials respectively. Constrained by available manufacturing processes and cost, design parameters are much less.

Constrained by available manufacturing processes and cost, design parameters are much less. In contrast, if manufacturing consideration is neglected, a container with various cross-section shapes are possible and it may be coated with heat insulation materials by AM. Here comes the question of how to stimulate designers who are novice in AM to think in an “AM way”.Following the similar concept of function-behavior-state design synthesis strategy, a function-behavior-functional entity-functional feature (FBFF) design strategy is proposed in this1937paper. Functional entity is defined as the abstraction of physical object to achieve the desired behavior. Functional feature is the concrete form of functional entity. Revisiting the case of the function “store water”, we can conclude that the desired behavior is “hold water from leaking”. The successive functional entity is “a semi-closed space” and the functional feature could be any surfaces satisfying volumetric requirements.

The successive functional entity is “a semi-closed space” and the functional feature could be any surfaces satisfying volumetric requirements. The mapping relations and knowledge input is shown in Figure 6. The mapping between function and behavior remains the same. Changes start at the mapping from behavior to functional entity with AM knowledge 1 (AMK1) as additional inputs to enrich options in the set of functional entity. AMK1 mainly refers to exceptional material properties enabled by AM, such as new materials and functional graded materials. The second change is the mapping from functional entity to functional feature with AM knowledge 2 (AMK2) as additional inputs to enrich functional feature alternatives. Physical phenomena refer to natural physiochemical laws such as Newton’s law and Bernoulli Equations. AMK2 mainly implies the innovative geometric, topology, and materials realized by AM. In this sense, functional features can be form features and material features.

AMK2 mainly implies the innovative geometric, topology, and materials realized by AM. In this sense, functional features can be form features and material features. For either a form feature or a material feature, it is modeled as a set of geometric, material distribution (topology), and material function. If the material is homogeneously distributed, the material function is a constant. If the material is function graded material (FGM), the material function could be represented by a voxel model [17] or an implicit model. It should be noted that the synthesis of functional features depends on other requirements as well if degrees of satisfaction of function [18] is considered.

It should be noted that the synthesis of functional features depends on other requirements as well if degrees of satisfaction of function [18] is considered. Compared with conventional design philosophy, as long as AMK1 and AMK2 are appropriately extracted, encoded, and delivered to designers at right sequence, FBFF design synthesis strategy can facilitate more design solutions to be found as shown in Figure 6 (the blue oval).A typical example of applications of the FBFF design synthesis approach is the morphing wing design. Traditional wings are designed as a compromise in geometry which allows the aircraft to fly at a range of flight conditions, but the flight performance at each condition is sub- optimal [19]. As one of the solutions, it requires the airfoil to change shape in different flight conditions. Assume that the desired behavior is “change skin profile with respect to speed”, then enabling AMK1 is “a micro-structure has multiple states”.

Assume that the desired behavior is “change skin profile with respect to speed”, then enabling AMK1 is “a micro-structure has multiple states”. The synthesized functional feature will be a surface with material which is sensitive to pressure and has multiple stable states. Due to the complicated internal structure, this synthesis design solution is promisingly realizable by AM.Set of functionsSet of behaviorSet of functional entitySet of functional featureRecognition abstractionPhysical phenomenaAMK1AMK2Figure 6 Function-behavior-functional entity-functional feature design synthesis strategy19385 Integrated FBFF in a redesign contextIn a redesign process, there are normally limited inputs such as a CAD model and possible design requirements.

Systematic quality management principles require complete tracking of product information from details back to design rationale and customer needs [20] and a mere geometric model is insufficient for recording such information. In this paper, the given CAD model serves as the starting point to understand working principles of the old design.

In this paper, the given CAD model serves as the starting point to understand working principles of the old design. The method can be taken is functional analysis, through which implicit functional information can be made explicit.CAD ModelCAD model of assemblyDesign requirementsFunction & function relation setFunction analysisFunction treeBehavior setFunctional entity setFunctional analysisC o n c e p t u a l d e s i g nFBFF methodDesign synthesisAMK3New functional feature setFunctional feature setFunction integrationNew design spaceFigure 7 Implicit information of a CAD model Figure 8 Redesign flowAs shown in Figure 7, FBFF-based functional information model should contain ontologies such as function & function relation, behavior, functional entity, and functional feature. Among these ontologies, only functional features are explicit in the CAD model.

Among these ontologies, only functional features are explicit in the CAD model. Once functional information becomes clear, FBFF method is applied to synthesize new functional features and these new functional features can be investigated if they can be consolidated into one design space. The criteria are AM knowledge 3 (AMK3). AMK3 should deal with four aspects of AM characteristics: part consolidation, embedded components, articulated kinematics, and general AM constraints (e.g. volume and tolerance). If selected functional features are suitable to be consolidated, then function integration process is implemented to generate a new design space as the new preliminary layout. Functional analysis, design synthesis, and function integration forms the conceptual design phase as shown in Figure 8.6 Case studyTo demonstrate how the proposed FBFF design synthesis approach works in the conceptual redesign of an assembly, a throttle pedal is exemplified with appropriate complexity.

As shown in Figure 9, the throttle pedal is comprised of 12 components without washers and fasteners. When a force is applied on the pedal, the lever transfers torque to the D-shape pin and the potential meter convert the angular motion to digital signal to control speed. To restrict the range of rotation, a rotation limit is needed. The maximum applied force is 250 lbs.1939Potential meterRotation limitbearingSealing ringRight caseshaftLeft caseTorsional springD-shape pinPedalLeverPinsFigure 9 exploded view of a throttle pedalControl accelerationCouple car bodyImport foot and import human E.External assembly interfaceConvert human E.

It should be noted that a concept of function relation is introduced to characterize relation between functions instead of physical relations. Function relations include decomposed-into, conditioned-by, enhanced-by, and external relations. Decomposed-into relation highlights function hierarchy while conditioned-by relation characterizes causal relations between functions. Enhanced-by relation signifies function redundancy. Because the system of interest is a part of a product, the system boundary is confined by external relations. As shown in Figure 10, functions and function relations construct the function tree. In some cases, some functions have already been integrated into one component.Design synthesisBased on the results of functional analysis, functional features should be identified and FBFF design synthesis should be conducted to generate new functional features.

However, due to the complexity of identifying all the functional features of the system, AMK3 is applied first. Part consolidation requires no relative motion, assembly access, and single material between two components; therefore, only pedal, pins, lever, right case, shaft, and D-shape pins are made as system of interest as depicted in Figure 11 a. As shown in Figure 11 b and c, functions are mapped to functional features. Due to the fact that the chosen system is only a portion of a bigger system, identification of system boundaries (SB) is primitive. In Figure 11 b, SB1 to SB6 all are mating interfaces while SB7 is for void consideration. In Figure 11 c, functions of components are decomposed into feature-level functions. For example, the hollow shaft needs to fulfill three functions: allow rotation, constrain rotation angle, and constrain DOF of torsion spring.

For example, the hollow shaft needs to fulfill three functions: allow rotation, constrain rotation angle, and constrain DOF of torsion spring. System boundaries sometimes need to deliver functions as well such as the feature “import human material” and the feature “transfer rotation”. In Figure 11 d, all necessary information is retrieved from the old design including system boundaries and functional features.In this case study, all the features except system boundaries have the potential to be redesigned. Taking the feature “allow rotation” and the feature “transfer torque” for instance, to realize function “allow rotation”, the functional entity is “a coaxial solid with the bearing”. Apparently, the shaft can be in any shape beyond SB1 as shown in Figure 11 e. In terms of function “transfer torque”, the functional entity is “a solid” and the functional feature is “a beam with one end in D shape”.

In terms of function “transfer torque”, the functional entity is “a solid” and the functional feature is “a beam with one end in D shape”. These modified functional features and the unmodified ones constitute the new functional feature set.Function integrationOnce the new functional feature set is complete, all the features in the set are checked for the possibility of function integration. If any two of them do not violate rules of “no relative motion”, “single material”, and “assembly access of a third part”, these features are endorsed as potential candidates. Afterwards, AM-specific rules can be applied to verify if the consolidated features violate manufacturing constraints. If not, these features form a new design space. Conceptual design of the new throttle pedal is finished.

If not, these features form a new design space. Conceptual design of the new throttle pedal is finished. It is worth mentioning that by using topology optimization to generate a reference model belongs to preliminary design to get the layout.1941SB1SB3SB4Transfer torqueSeparate solid from dustAllow rotationConstrain DOF of springSB2Amplify torqueSB7SB5Constrain rotation angleDetailed view of the feature “constrain DOF of spring”SB6Import human materialDetailed view of the feature “transfer torque”(a) System of interest(b) identify system boundary(c) identify functional features(d) old functional feature set(e) modified functional features(f) New design spaceFigure 11 design synthesis and function integration process of the throttle pedal7 Feature researchAM facilitates much more design freedom than conventional manufacturing processes did, while how to take full advantage of these capabilities have not paced up with the advances.

This paper focuses on solving the problem of how to redesign an assembly at conceptual design stage in the context of AM in order to stimulate novice AM users to achieve incremental innovation. In the first place, popular design methodologies working at conceptual design stage are reviewed. These design methodologies are in different representations, but their ways of generating a design concept can be concluded as functional analysis and design synthesis. In the second place, the impact of AM-enabled design freedom on functional analysis and design synthesis are studied respectively. AM has minor effect on functional analysis process. In contrast, design synthesis stage is analyzed by using function-behavior-state functional modeling method and evidence proves that AM-enabled design freedom enriches the mapping relations between the desired behavior and the corresponding physical structures.

In the third place, function-behavior- functional entity-functional feature (FBFF) design synthesis method is proposed to show how, when, and what AM knowledge should be applied in order to facilitate product innovation. In the end, FBFF method is applied in a redesign context and proved that it is able to embrace AM- enabled function integration.In the future, two research directions need to be explored more. The first one is that conceptual design stage is so function-immersed and the unsolved issues with function description makes it difficult to encode functional information in a uniform way. The other one is how to enrich the mapping from functional entity to functional features in the proposed FBFF design1942synthesis method. There are two possible ways to enhance creativity. One possible way is to do analogical functional reasoning with existing products.

There are two possible ways to enhance creativity. One possible way is to do analogical functional reasoning with existing products. As an example, Oregon State University managed to build a repository with more than 6000 artifacts and relate functions with various parts [21]. The other way is to establish an innovative design database serving as inspiration for new feature design [22].ACKNOWLEDGEMENTFinancial support from the National Sciences and Engineering Research Council of Canada Discovery Grant RGPIN 436055-2013 and McGill Engineering Doctoral Award (MEDA) is acknowledged with gratitude.Reference[1] ASTM, 2012, "F2792. 2012. Standard Terminology for Additive Manufacturing Technologies," ASTM F2792-10e1, ASTM International. [2] Gibson, I., Rosen, D. W., and Stucker, B., 2010, Additive Manufacturing Technologies:Rapid Prototyping to Direct Digital Manufacturing, Springer US.

[2] Gibson, I., Rosen, D. W., and Stucker, B., 2010, Additive Manufacturing Technologies:Rapid Prototyping to Direct Digital Manufacturing, Springer US. [3] Yang, S., and Zhao, Y., 2015, "Additive manufacturing-enabled design theory and methodology: a critical review," The International Journal of Advanced Manufacturing Technology, pp. 1-16. [4] LAVERNE, F., SEGONDS, F., ANWER, N., and Marc, L., 2015, "Assembly-based methods to support product innovation in Design for Additive Manufacturing: An exploratory case study," Journal of Mechanical Design. [5] Suh, N. P., 1990, The principles of design., Oxford University Press, New York, USA. [6] Rizzuti, S., "The Importance of Teaching Functional Analysis in Product Design Courses," Proc. DS 69: Proceedings of E&PDE 2011, the 13th International Conference on Engineering and Product Design Education, London, UK, 08.-09.09. 2011.

[11] Umeda, Y., Ishii, M., Yoshioka, M., Shimomura, Y., and Tomiyama, T., 1996, "Supporting conceptual design based on the function-behavior-state modeler," Ai Edam, 10(4), pp. 275-288. [12] Tomiyama, T., Gu, P., Jin, Y., Lutters, D., Kind, C., and Kimura, F., 2009, "Design methodologies: Industrial and educational applications," CIRP Annals-Manufacturing Technology, 58(2), pp. 543-565. [13] Booth, J. W., Reid, T. N., Eckert, C., and Ramani, K., 2015, "Comparing functional analysis methods for product dissection tasks," Journal of Mechanical Design, 137(8), p. 081101. [14] Culverhouse, P., 1995, "Constraining designers and their CAD tools," Design Studies, 16(1), pp. 81-101. [15] Tomiyama, T., Umeda, Y., Ishii, M., Yoshioka, M., and Kiriyama, T., 1996, "Knowledge systematization for a knowledge intensive engineering framework," Knowledge intensive CAD, Springer, pp. 33-52.1943[16] Umeda, Y., Takeda, H., and Tomiyama , T., "Function, behavior and structure.," Proc.

J., and Mäntylä, M., 1995, Parametric and feature-based CAD/CAM: concepts, techniques, and applications, John Wiley & Sons. 2006, [21] http://function2.mime.oregonstate.edu:8080/view/index.jsp. [22] Bin Maidin, S., 2011, "Development of a design feature database to support design for additive manufacturing (DfAM),"Dissertation, Loughborough University, UK.University,O.S.,"Designrepository,"1944.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceConditional generative adversarial networks for in-situ layerwise additive manufacturing dataChristian Goberta, Edel Arrietab, Adrian Belmontesb, Ryan B. Wickerb, Francisco Medinab, Brandon McWilliamsca Oak Ridge Associated Universities, Belcamp, MD b The University of Texas at El Paso, El Paso, TX c CCDC Army Research Laboratory, Aberdeen Proving Ground, MD AbstractConditional generative adversarial networks (CGANs) learn a mapping from conditional input to observed image and perform tasks in image generation, manipulation and translation. In-situ monitoring uses sensors to obtain real-time information of additive manufacturing (AM) processes that relate to process stability and part quality.

In-situ monitoring uses sensors to obtain real-time information of additive manufacturing (AM) processes that relate to process stability and part quality. Understanding the correlations between process inputs and in-situ process signatures through machine learning can enable experimental-driven predictions of future process inputs. In this research, in-situ data obtained during a metallic powder bed fusion AM process is mapped with a CGAN. A single build of two turbine blades is monitored using EOSTATE Exposure OT, a near-infrared optical tomography system of the EOS M290 system. Layerwise images generated from the in-situ monitoring system were paired with a conditional image that labeled the specimen cross-section, laser-scan stripe overlap and z-distance to part surfaces.

Layerwise images generated from the in-situ monitoring system were paired with a conditional image that labeled the specimen cross-section, laser-scan stripe overlap and z-distance to part surfaces. A CGAN was trained using the turbine blade data set and employed to generate new in-situ layerwise images for unseen conditional inputs.Keywords: machine learning, additive manufacturing, conditional generative adversarial networks, in-situ monitoringIntroductionMetallic powder bed fusion (PBF) additive manufacturing (AM) is a process in which near-net shape parts are built directly from digital files in a layer-by-layer manner [1-4]. In-situ monitoring uses sensors to obtain real-time information during the build, aiming to build correlations between in-situ process signatures and post-process material quality [5-7].

In-situ monitoring uses sensors to obtain real-time information during the build, aiming to build correlations between in-situ process signatures and post-process material quality [5-7]. The PBF AM EOS M290 systems employ a high-resolution camera at high frequency in the near-infrared range, capturing a layer-wise image representative of the material melting behavior across the build-plate [8]. The correlations between post-process material quality and laser-material interactions is understood to be impactful [1-7]. Building an understanding through machine learning tools between in-situ information and process inputs is therefore desirable.Machine learning encompasses classes of computer algorithms that map inputs to desired outputs by optimizing parameters without explicitly being programmed [9]. Convolutional neural networks (CNNs) are multi-stage algorithms that can learn features through application of trainable convolutional filters [10-12].

Convolutional neural networks (CNNs) are multi-stage algorithms that can learn features through application of trainable convolutional filters [10-12]. Conditional generative adversarial networks (CGANs) attempt to learn a mapping from a conditional input to an observed images by pitting two neural networks, a generator and discriminator, against each other. The generator attempts to produce realistic-looking images from conditional inputs while the discriminator aims to distinguish192between generated images (i.e. fake) and observed images (i.e. real) [13-15]. The pix2pix [14] network is a demonstrated CGAN for images up to a resolution 256×256 pixels, using a U-Net [16] architecture as the generator and a patch-based CNN as the discriminator.In this research the pix2pix network architecture was applied to EOSTATE Exposure OT in-situ data generated from two printed turbine blades. High-resolution near-infrared images were generated for each layer of the build.

High-resolution near-infrared images were generated for each layer of the build. Identical sized conditional images were constructed which contained information on the STL cross-section, laser-scan stripe overlap, and z-distance to part surfaces. A subset of the data was used to train the pix2pix network, the remaining data was used for validation. Conditional inputs were generated from outside the dataset using geometric shapes to view the CGAN’s generalization of new process inputs.Material and MethodsSpecimens and ProcessTwo turbine blades were printed side by side on an EOS M290 system. Each turbine blade occupied a bounding box of 80mm height, 62.5mm width and 25.5mm depth, a 3D view is depicted in Figure 1. Process parameters were set for Inconel 718; scan strategy was stripe-based with a width of 10mm, overlap of 0.12mm, and hatch distance of 0.11mm while the laser was set to 285W at 960mm/s; contours are offset by 0.012mm.

The upskin process used 153W at 600mm/s with hatch distance of 0.09mm, and the downskin process used 95W at 785mm/s with hatch distance of 0.1mm. Both, upskin and downskin consisted of being 3 layers in thickness. Layer height was set to 0.04mm with 2000 build layers.Figure 1. 3D view of single turbine blade.In-situ Data StreamFor each build layer, the EOS M290 records a near-infrared high-resolution optical image of the build surface recording the energy application in the melt-pool [8]. The sCMOS camera detects near-infrared images (900nm) by using a bandpass filter to block any emissions from plasma and reflected light from the laser. The Exposure OT system measures the intensity of the emissivity of the molten material, in a long-exposure photograph of the build plate at every layer.

The Exposure OT system measures the intensity of the emissivity of the molten material, in a long-exposure photograph of the build plate at every layer. The amount193of emissivity can be interpreted as a fraction of the residual energy from the melting process; moreover, the law of thermal radiation formulated by Max Planck correlates emitted energy to temperature according to wavelength. In this context, image zones with marked intensity variations may be indicators for flaws and defects; it is important to mention, that a variation in intensity does not necessarily imply defects all the time. For example, too much emissivity could be detected in narrow geometries, given the saturation of energy due the short scan vectors, which may result in distortions or local metal vaporization. On the other hand, it is expected to have higher emissivity detected from the laser-scan stripe overlaps given that the laser spends more time along them.

On the other hand, it is expected to have higher emissivity detected from the laser-scan stripe overlaps given that the laser spends more time along them. In this case, high emissivity from laser-scan stripe overlaps rarely represent defects. The resolution of the layerwise images is approximately ~125 μm/pixel and are 8-bit RGB images where increasing image intensity corresponds to higher emissivity and increasing energy application. A sample image from the monitoring software is displayed in Figure 2, showing the captured image from a single layer of the turbine blade build.Figure 2. Image of in-situ data stream from Exposure OT of single build layer of two turbine blades.In figure 2 the stripes of red can clearly be seen that are a result of the laser-scan stripe overlaps. The dark red regions are a result of the upskin and downskin process parameters that are utilized for near-part surface regions.

The dark red regions are a result of the upskin and downskin process parameters that are utilized for near-part surface regions. The increase in image intensity can be seen in the regions of acute angles created between the part contours and laser-scan stripe overlaps, most-likely due to the loitering of the laser in small regions during scanning.Data SetThe data set for the CGAN network comprises of paired observed in-situ images and corresponding conditional images for each sample. A sample was generated for each turbine blade and build layer. The in-situ images were 8-bit RGB 512×512 pixels images cropped to the center of each turbine blade. The conditional images were 8-bit RGB 512×512 pixel images, centered on the turbine blade as well, with each image channel representing a process input. The red channel was used to indicate the locations of the part obtained from the STL, pixels with a red channel value equal to 255 indicated a part region.

The red channel was used to indicate the locations of the part obtained from the STL, pixels with a red channel value equal to 255 indicated a part region. The green channel was used to indicate the location of laser-scan stripe overlaps, pixels with a green channel value equal to 255 indicated a region of laser-scan stripe overlap. The blue channel was used to indicate the distance in layers for an upskin or downskin process, pixels were assigned a standardized value that was mapped to the RGB space. Pixels were given an integer value from -3 to 3 standardized to [0 255] that indicated a pixels distance related to the upskin and downskin parameters. An input value of zero (standardized to194127 in RGB space) represented no upskin or downskin parameters. A pair of observed and conditional images is displayed in Figure 3.Figure 3.

An input value of zero (standardized to194127 in RGB space) represented no upskin or downskin parameters. A pair of observed and conditional images is displayed in Figure 3.Figure 3. Sample of observed in-situ image (left) and corresponding conditional image (right).For network training, images were resized to 256×256 pixels due the limitations in resolution for the pix2pix methodology and RGB values were mapped from [0 255] to [0 1]. Generating the conditional images is time-consuming and only 582 samples were used, a sample encompasses a pair of observed and conditional images. In Isola et al. [14], demonstrations of the pix2pix network had been shown with data sets containing only 400 images.Machine Learning NetworkCGANs contain two neural networks, a generator (i.e. to generate realistic looking samples given a conditional input) and a discriminator (i.e. discriminate between observed and fake images).

to generate realistic looking samples given a conditional input) and a discriminator (i.e. discriminate between observed and fake images). The generator input and output are the conditional image and the observed image, respectively. The generator is defined by down-sampling and up-sampling modules. A down-sampling module comprises of n number of 4×4 convolutions of stride 2, leaky ReLU activation [17] and batch normalization [18]. An up-sampling module comprises of a 2×2 up-sampling, n number of 4×4 convolutions of stride 1, leaky ReLU activation and batch normalization. The end of an up- sampling module is where the U-net skip connections are connected to the corresponding output of a down-sampling module. The generator comprises of five down- and up-sampling modules, from 256×256 pixels to 8×8 pixels. The number of filters used in each module increases with decreasing image size; being 32, 64, 128, 256, and 512 filters for each of the corresponding down- and up-sampling modules.

The number of filters used in each module increases with decreasing image size; being 32, 64, 128, 256, and 512 filters for each of the corresponding down- and up-sampling modules. The last layer of the generator is a 3-channel output to match the RGB observed images, using 32 4×4 convolutions and leaky ReLU activation followed by 3 4×4 convolutions with sigmoid activation [17].The discriminator input is a channel-wise concatenated array of the conditional input and observed image (or generated fake image depending on the sample). The output of the discriminator CNN is a patch multi-channel image of 8×8 pixels. For real and fake samples the output patches are set to a value of 1 and 0, respectively. The discriminator CNN includes five down-sampling modules with the image size reduced from 256×256 pixels to 8×8 pixels. Each down-sampling module195comprises of n number of 4×4 convolutions with stride of 2, leaky ReLU activation and batch normalization.

Each down-sampling module195comprises of n number of 4×4 convolutions with stride of 2, leaky ReLU activation and batch normalization. The final layer of the discriminator CNN is an 8×8 output generated from a 3×3 convolutional filter with sigmoid activation. The number of filters used in each module increases with decreasing image size; being 32, 64, 128, 256, and 512 filters for each of the corresponding down-sampling modules.Network Training, Validation and AnalysisThe set of turbine blade samples are divided into a training and validation set at a ratio of 80/20, randomly sampled from the set. The training data was used to update network parameters. The validation data was only used to evaluate the network performance at the end of a training epoch and in the post-process analysis.

The training data was used to update network parameters. The validation data was only used to evaluate the network performance at the end of a training epoch and in the post-process analysis. Training a CGAN involves the training of the discriminator and generator.During training the discriminator attempts to learn parameters that can enable it to discriminate between real and fake images by minimizing the loss function. The loss function for the discriminator was measured through the mean squared error of the patches and accuracy was used as a metric to gauge performance. For each iteration the discriminator is trained, the generator is also trained. The generator is trained to maximize the loss of the discriminator for samples that are fake (i.e. fool the discriminator). In this research, a batch size of one was used and the Adam optimizer [19] with learning rate of 0.0001, beta one of 0.9 and beta two of 0.999.

fool the discriminator). In this research, a batch size of one was used and the Adam optimizer [19] with learning rate of 0.0001, beta one of 0.9 and beta two of 0.999. The training was completed for 200 epochs, where each epoch included a number (equal to the number of training samples) of randomly sampled training iterations.Results and DiscussionNetwork Training and ResultsThe CGAN was trained for 200 epochs, with the validation loss of the discriminator and generator being recorded for each epoch, displayed in Figure 4. Figure 5 displays the validation accuracy of the discriminator for each epoch.1965- Discriminator - -Generator4.5~ 4 ~ ~ 3.5 wi 3 & 2.5 m:,C2~1.50.5 ~-~--~=~-~~-=-~~~~~~~--~-~ 100 200 Epoch020406080120140160180Figure 4.

Validation loss of the generator (orange) and discriminator (blue) from each training epoch.0.5''''0.450.40.35~~ ~N~\ ~ ~lj/\0.3c; ~ 5 0.25 0 <C0.2 -0.1 50.10.05\.!0020406080100 Epoch120140160180200Figure 5. Validation accuracy of the discriminator from each training epoch.Examining Figure 5, the lowest accuracy was achieved around epoch 26. Figure 6 displays a set of samples generated from the generator network at epoch 26.197Figure 6. Validation samples and generated images from the epoch 26 network. The top row contains observed images, middle row contains conditional images, and the bottom row contains generated images, paired column-wise.The generator tended to represent the high emissivity in the acute angles formed by the laser-scan strip overlap regions and STL contours. The generator seems to match the high emissivity on and around laser-scan stripe overlaps.

The generator seems to match the high emissivity on and around laser-scan stripe overlaps. There is an inability to represent the noise seen in ground truth images, due to the underlying architecture of the network.Figure 7. Sample of turbine with observed image (left), conditional image (center) and generated image (right).Figure 7 shows an example of how the network tended to overestimate the average emissivity intensity is some circumstances. The assumption for this phenomenon is first that the network does not have a complete understanding of the part geometry (i.e. layers below) and subsequent impact198that would have on energy saturation. Second the distribution of data was drawn from a section of the turbine blade with larger emissivity values, possibly biasing the generator.

Second the distribution of data was drawn from a section of the turbine blade with larger emissivity values, possibly biasing the generator. Third the network does not know the true path of the laser in the regions of the stripes and the sequence of neighboring stripes which would have an impact on energy saturation.Generated Novel SamplesWith a trained generator, novel inputs can be processed and studied. In Figure 8, three unique geometries with two laser-scan stripe overlap spacings are displayed, generated from the trained network of epoch 26.Figure 8. Generate novel samples. There are three unique geometries (columns) and two unique laser-scan stripe overlap spacings (rows).Figure 8 shows an example of how the average intensity of the specimen slice decreases with increasing separation of laser-scan stripe overlaps, due to less stripe overlaps over the entire layer.

In Figure 9 two geometries are depicted with two instances of laser-scan stripes overlap regions.199Figure 9. Generated images for two unique geometries (rows) and two unique laser-scan stripe overlap regions (columns)In Figure 9 the geometry of the circle is invariant to the angle of the laser-scan stripe overlap, unlike in the square, which impacts the average intensity of the image as a result of the acute angles created between scan-line stripe overlaps and part contours.ConclusionA CGAN was demonstrated for in-situ data generated using the EOSTATE Exposure OT system on the EOS M290. The trained generator was able to mimic characteristics of the in-situ images, but failed to represent the stochastic nature of the images and in instances overestimated the average emissivity. This work demonstrated an initial intuition of CGANs with in-situ data. Future research needs to explore the generalization capabilities of CGANs, with increasing scope.

This work demonstrated an initial intuition of CGANs with in-situ data. Future research needs to explore the generalization capabilities of CGANs, with increasing scope. Ideally the conditional pair to an observed image would encompass the complete PBF AM process input space, i.e. geometry, material, process parameters, with the aim of simulating PBF AM builds through CGANs. Applying CGAN architectures that can accommodate larger resolution images would be utilized as well, as the limit to this approach was 256 pixels and ~250 microns/pixel. Once a CGAN is trained, generating images is relatively time inexpensive. Ideally, simulating builds with CGANS would complement the slow and computational expensive physics based models with data-driven predictions and aid in optimization of process inputs for PBF builds.References[1] Frazier, W Metal additive manufacturing: a review, Journal of Materials Engineering andPerformance, 23:1917–1928, April 2014[2] I. Gibson, D. Rosen, B.

Gibson, D. Rosen, B. Stucker, Additive manufacturing technologies: rapid prototyping todirect digital manufacturing, Springer, New York, 2009.200[3] Sames W J, List, F A, Pannala S, Dehoff R R, Babu S S 2016 The metallurgy and processing science of metal additive manufacturing, International Materials Review, 61:315-360 March 2016[4] Debroy T, Wei H L, Zuback J S, Mukherjee T, Elmer J W, Milewski J O, et al. Additive manufacturing of metallic components – process, structure and properties, Progress in Materials Science, 92:112-224, October 2017[5] Everton SK, Hirsch M, Stravroulakis P, Leach RK, Clare AT, Review of in-situ process monitoring and in-situ metrology for metal additive manufacturing, Materials & Design, 95:431-435, April 2016.[6] G. Tapia, A.

Elwany, A review on process monitoring and control in metal-based additive manufacturing, Journal of Manufacturing Science and Engineering, 136(6):60801–60811, December 2014[7] Zhang B, Li Y, Bai Q, Defect formation mechanisms in selective laser melting: a review,Chinese Journal of Mechanical Engineering 30:515-527, March 2017[8] https://www.eos.info/press/case_studies/optical-tomography-enables-considerable-savings-in-quality-assurance[9] Bishop C M, Pattern Recognition and Machine Learning, Springer, New York, 2006 [10] LeCun Y, Bengio Y, Hinton G, Deep learning, Nature 521:436-444, May 2015 [11] LeCun Y, Kavukcuoglu K, Farabet C, Convolutional networks and applications in vision, Proceedings of IEEE International Symposium on Circuits and Systems, 253-256, 2010 [12] Krizhevsky A, Sutskever I, Hinton G E, ImageNet classification with deep convolutionalneural networks Advances in Neural Information Processing Systems, 25 , 2012[13] Goodfellow I, Pouget-Abadie J, Mirza M, Xu B, Warde-Farley D, Ozair S, Courville A, Begio Y, Generative adversarial networks, Advances in Neural Information Processing Systems, 27, 2014[14] Isola P, Zhu J, Zhou T, Efros A, Image-to image translation with conditional adversarial networks, Proceedings of the IEEE conference on computer vision and pattern recognition, 1125-1134, 2017[15] Wang T, Liu M, Zhu J, Tao A, Kautz J, Cantanzaro B, High-resolution image synthesis and semantic manipulation with conditional GANS, Proceedings of the IEEE Conference on Computer Vision and Pattern Recognition, 2018[16] O.

Ronneberger, P Fischer and T. Brox, U-Net: Convolutional Networks for Biomedical Image Segmentation, International Conference on Medical image computing and computer- assisted intervention, 234-241, October 2015[17] Nwankpa C, Ijomah W, Gachagan A, Marshall S, Activation functions: comparison of trendsin practive and research for deep learning , November 2018[18] Ioffe S, Szegedy C, Batch normalization accelerating deep network training by reducinginternal covariate shift, May 2015[19] Kingma D, Lei Ba J, Adam: a method for stochastic optimization, January 2017201.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCONFIGURATION CONTROL FOR ADDITIVE MANUFACTURING DIGITAL TWINSD. W. Gibbons¹, and H. Ko² ¹Department of Industrial Engineering, Stellenbosch University, Stellenbosch, 7600, RSA²School of Manufacturing Systems and Networks, Arizona State University, Mesa, AZ 85212, USAAbstractThe additive manufacturing workflow is a ductile entity, often varying depending on the design, the product, the process, the material, and the application. Information models and schemas have been developed that can provide structure to data and information throughout the workflow. The result has been a well-characterized outline of an additive manufacturing digital thread.

The result has been a well-characterized outline of an additive manufacturing digital thread. However, implementation-specific details are often missing from these characterizations, creating challenges in establishing part-specific workflows necessary for product configuration control and management. While software vendors are increasingly filling this gap, a software-agnostic workflow is yet to be defined. This paper investigates standardized, the additive manufacturing workflow and establishes configuration-control approach including formats and interoperability while addressing versioning, digital rights, and ownership.thefundamentalsofaIntroductionAs industry moves towards the fourth industrial revolution and digital supply chains, the need to establish transparency and accountability across organizations has only grown.

Increasingly sophisticated manufacturing processes have led to heightened vigilance and abundance of caution when introducing new parts into a supply chain. The digitalized nature of these processes, however, has created new opportunities to leverage data throughout the product lifecycle for reasons including traceability, improved efficiencies, compliance and innovation [1]. These opportunities are nowhere more apparent in additive manufacturing (AM), an inherently digital production process.thanAM machines fabricate digitized designs using a that follows a digitally planned production path. Given the intricate process complexities, additively manufactured parts are often considered “unique” when compared to their traditionally manufactured counterparts. This uniqueness holds true for the digitalization characteristics of the processes as well, in that the extent of the AM digital workflow sets itself apart from other manufacturing workflows.

This uniqueness holds true for the digitalization characteristics of the processes as well, in that the extent of the AM digital workflow sets itself apart from other manufacturing workflows. Here, we will call the connection of digital components throughout this workflow the AM digital thread [2]. We will call aggregated data associated with a single instance of a part traversing this thread a digital twin, with the twin “growing” in size from the generation of new data at every layer.layer-by-layer processThe AM digital thread incites diverse transformations in the digital twin when progressing from a raw design to a qualified product. In transitioning from a design to a manufactured part, the AM digital thread can be defined by several stages including CAD Design, Tessellated Geometry, Sliced Geometry, Build File, Part with Build Data, and Part with Evaluation Data. Each of these stages will introduce new information requirements to be associated with the digital twin.

Each of these stages will introduce new information requirements to be associated with the digital twin. These digital twins provide a snapshot, or representation, of the state of an AM part at any point in time along the workflow path. Managed correctly, these snapshots can provide vital insights into the state of a part, from design to fabrication to inspection.In quality control for AM, establishing digital provenance has become a necessary part of the part-qualification process. When accepting a final part, it is important to have confidence in the processes used in the fabrication of the part, since differences in implementation can result in differences in part quality. Two parts progressing through the same thread can lead to two different digital twins, even if the process configuration and materials269are similar. By providing each part with a unique “signature”, the digital twin provides an important resource from which this confidence can be gained.

By providing each part with a unique “signature”, the digital twin provides an important resource from which this confidence can be gained. However, the underlying data has its own uncertainty even if it is not well-defined and not well-understood. As a consequence, the digital thread has an uncertainty, which is a combination of these individual uncertainties. When relying on a digital thread to guide the provenance of part, not understanding the context and fidelity of the associated information limits the amount of insight that can be attained.This paper investigates configuration management and control of data throughout the AM product lifecycle. A data model is presented that defines 1) how critical configuration data throughout this lifecycle can be identified, 2) how the associated data formats can be packaged, 3) how to transform these data packages from one configuration to the next and 4) how the information gains or losses transition through this workflow.

This data model provides an approach to identify and control the critical data along the product lifecycle, thereby providing a workflow reference for future AM product-realization endeavors and an architecture for implementing AM digital twins and threads. The model aims at being process and industry agnostic. It also does not dictate which approach is optimal but provides the user with the needed information to make the correct choices and identify the required configuration items for their particular application.Background2.1. Additive Manufacturing as a Digital ProcessAM is seen as an inherently digital production process. Kim, Witherell, Lipman, and Feng [2] refer to this concept that includes the magnitudes of associated information, formats and software as the “digital spectrum”. For a given workflow within this spectrum, the digital thread links the path of the information that is gathered throughout the lifecycle [2].

For a given workflow within this spectrum, the digital thread links the path of the information that is gathered throughout the lifecycle [2]. Understanding and managing this thread is vital to truly take advantage of afforded opportunities and insights into a part and process. Mies, Marsden, and Warde [1] note the importance of infrastructure for data capture, availability, and discovery and how this impacts intellectual property, the tools for data mining, data federation and integration, quality standards and best practices. They further noted the challenges this poses since it requires cultural shifts from traditional manufacturing approaches. Mies et al. [1] coin this broader concept “AM informatics”.

They further noted the challenges this poses since it requires cultural shifts from traditional manufacturing approaches. Mies et al. [1] coin this broader concept “AM informatics”. Ultimately, the digital thread provides an improved vision into the product lifecycle; this, in turn, benefits complex production processes such as AM.Kim, Witherell, Lu and Feng [3] state that maintaining consistency across AM builds is a challenge, noting that variations in the process, material, or geometry cause such variability between builds. Further contributing factors to such variability include:• • •The use of different data formats and schemas. Proprietary file formats and software. Closed AM machines leading to inconsistencies between AM machine models.

Proprietary file formats and software. Closed AM machines leading to inconsistencies between AM machine models. Lack of a standard for defining the critical data throughout the AM product lifecycle required to reliably reproduce AM parts.Each of these factors compounds the challenges of creating and maintaining a digital twin that can be consistently accessed and maintained by various stakeholders. This is particularly true when stakeholders are dispersed across an organization or organizations. This is often the case with an AM part since this part is a component of a larger product made by a distributed supply chain. This distributed manufacturing capability is often cited as an advantage of the AM process itself, whereby a component can be designed at a head office and sent out to the locations where the product must be e manufactured on-site. Doing so reduces lead times and the need for large, on-site inventories.

Doing so reduces lead times and the need for large, on-site inventories. This concept of distributed manufacturing can be extended in cases where certain post- processing or component testing and qualification are performed at different locations or by different organisations. Not having a defined and standardized approach for characterizing the AM digital twin poses many challenges for situations such as these just described.270The need for a standardized approach that addresses representations across the AM product lifecycle is identified as a current gap [4], though the AM workflow and digital thread have been characterized previously. Previously published literature on the topic include Nassar and Reutzel [5] who proposed a digital thread for AM, Kim et al. [2] who investigated the information requirements workflow of the AM digital spectrum, Mies et al. [1] provided an overview of AM informatics and how this drives the digital thread, Kim et al.

[2] who investigated the information requirements workflow of the AM digital spectrum, Mies et al. [1] provided an overview of AM informatics and how this drives the digital thread, Kim et al. [3] proposed a tiered AM information map including data models and key attributes that span from design through to qualification, and Bonnard, Hascoët, Mognol and Stroud [6] presented a STEP-NC compliant data model for an AM digital thread. While these efforts help establish an accepted flow of information, they did not aim to address how the flow impacts the establishment, or configuration, of the digital twin.2.2. Configuration Management & the Need Thereof in Additive Manufacturing ProductionConfiguration-management techniques are required to control and manage data within AM production.

Configuration Management & the Need Thereof in Additive Manufacturing ProductionConfiguration-management techniques are required to control and manage data within AM production. Configuration is defined as “a collection of an item’s descriptive and governing characteristics that can be expressed in functional and physical terms, this represents the requirements, architecture, design and implementation that define the version of the system and its components” [7]. Whereas configuration management is defined as “a process for establishing and maintaining consistency of a product’s performance, functional and physical attributes with its requirements, design and operational information throughout its life” [7]. Configuration-management standards and handbooks such as SAE EIA-649, ISO 10007 and MIL-HDBK-61B define the configuration management process in five key functions.

Configuration-management standards and handbooks such as SAE EIA-649, ISO 10007 and MIL-HDBK-61B define the configuration management process in five key functions. These functions are defined in green in Figure 1 per the MIL-HDBK-61B definition, and subsequently used as guidelines to discuss the adoption of configuration management for AM in subsections of 2.2.1 to 2.2.62.2.1. Configuration planning and managementConfiguration management for AM should be planned during the earliest project stage. Procedures should be developed for managing the configuration of the design and relevant data through to its disposal. Additional considerations should be made for configuration between organizations to ensure each organisation’s approach and coding schemes are consistent and compatible.2.2.2.

Additional considerations should be made for configuration between organizations to ensure each organisation’s approach and coding schemes are consistent and compatible.2.2.2. Configuration identificationIdentification of Configuration Items (CI) requires the knowledge of which items have an effect on the integrity of the component data as well as which data items are required for compliance consistency. All CIs and associated AM formats should be assigned an identifier and revision controlled.2.2.3. Configuration control (Process required to change a CI and re-baseline it)Configuration control can be defined as “a systematic process that ensures that changes to released configuration documentation are properly identified, documented, evaluated for impact, approved by an appropriate level of authority, incorporated, and verified” [8].

In documenting a CI change, critical information to be captured includes the nature of the configuration change, the identification of the previous state and the current state of the configuration item, and information of the data transformation and losses.2.2.4. Configuration status accounting (Traceability)The configuration data shall be stored in a system that allows for retrieval of configuration statuses. [8]. This allows for the component’s full digital twin to be accessed when required anywhere along its digital thread.2.2.5. Configuration verification and auditConfiguration baselines need to be periodically audited to verify their contents and ensure conformance. This involves a functional and physical verification of the component configuration.271Figure 1: DoD configuration management activity model [8].2.2.6.

This involves a functional and physical verification of the component configuration.271Figure 1: DoD configuration management activity model [8].2.2.6. Conceptualizing Configuration Management for AM ProductionConfiguration management is a process that is begun early in the product design phase and aims at ensuring that the design intent is realized by monitoring and controlling the subsequent processes. In digital production processes such as AM, applying configuration management practices requires the configuration of both the physical aspects and their associated digital twins. The resulting digital twins and corresponding digital thread must accurately model their real-world counterparts, leading to full digital traceability and record of the AM product.Focusing on the AM production digital twin, the application of configuration management must aim at reducing impermissible and unintended changes and monitor/record the permissible changes.

Such changes could be increasing the component wall thicknesses in the CAD digital twin to compensate for material removal in post- processing stages, converting the CAD digital twin into an AM machine-readable format or even the orientation and location of the design in the build envelope. These are examples of changes that may have an effect on the design intent of the component and therefore such configurations should be managed. It is not always the case that subsequent processes require each of the preceding configurations, but it is still necessary to maintain configuration control. This concept is presented in Figure 2, where the focus is on what design and data configuration changes are permitted if the final product is to fulfill its original intent, and how should these be defined and controlled.Figure 2: Conceptual diagram of data configuration within AM workflow.2722.3.

Towards a Reference Data ModelThe success of any digital twin depends on its accuracy when compared to the physical reality it aims to represent. The more real-world data of high fidelity that can be represented by the digital twin the greater its accuracy. The application of digital-twin concepts to AM has its benefits and challenges. The key benefit being that since the AM process is inherently a digital one, the AM process is an excellent candidate for the application of digital twin concepts when compared to some of the more traditional manufacturing processes. Conversely, challenges are created when building digital twins from the large amount, diversity, and complexity of the variety of data involved during the AM process and the product lifecycle as a whole.As a part transitions through its AM digital thread, it is configured in various ways that provide benefit for the activities being performed at the relevant stages throughout the workflow.

There is a need for a definition of:The stages in such AM workflows and the relevant product configurations. ● What the data inputs and outputs are exhibited at these stages. ● How the data can be formatted and configured at each of these stages to ensure high fidelity.A reference data model that addresses these needs provides benefits for planning and managing AM production. Such a reference data model can be used for defining configuration baselines and, in turn, reduces the risk of unintentionally changing data throughout the workflow. Establishing a reference data model does not have to be a universal approach, but instead based on the workflow of a specific organization or a subset of specific stages from the digital thread. The creation of a data package is one way to down-select from the digital thread.Figure 3 illustrates the relationship between the digital thread and the digital twin.

The creation of a data package is one way to down-select from the digital thread.Figure 3 illustrates the relationship between the digital thread and the digital twin. Note that the arrows in Figure 3 indicate that the data package facilitates the selection of data requirements for the digital twin from the digital thread. Left uncontrolled, these relationships pose various challenges, the main being determining what data is required in the data packages that are distributed between the locations or organisations and how is this data to be configured in order for the design to be manufactured and evaluated to fulfil its design intent. Here, configuration management establishes consistency in how the data package requirements are met.The next section will adopt the concepts put forth in Section 2. The proposed method aims to integrate established data management concepts with specific AM needs.

The proposed method aims to integrate established data management concepts with specific AM needs. A result is a conceptual approach to establishing configuration management practices for AM.Figure 3. Configuration Management establishes consistency in HOW information requirements are met.273ApproachAs alluded to in the previous sections, configuration management is not established as a “one size fits all”. Instead, configuration management can be thought of as providing a “filter” when establishing a digital twin, as the plan:Defines the allowable representations on which requirements can be met. ● Identifies allowable formats and configurations. ● Establishes consistency in how data is captured and represented. ● Essential for consistent evaluation and qualification of families of AM parts.In establishing a data reference model, the realization of the configuration management plan is based on the data package requirements and the organizational workflows.

As such, configuration management requires a well- defined approach when establishing digital twin requirements. Such an approach is presented as four steps in Figure 4, where the steps are, 1) identifying WHAT information from the “digital spectrum” is leveraged by the workflow and used to establish the digital thread, 2) identifying HOW this information is represented through available format types, 3) establishing specific data requirements based on data package requirements, and 4) mapping data package requirements to appropriate data formats. The steps are designed in such a way that steps one and two may be developed for a general process, while steps three and four are likely to be specific to a given scenario.Figure 4: The approach for establishing configuration management for AM production.The approach is expanded upon in the following subsections:1. Establish the AM workflow. 2. Identify data types and requirements. 3.

Establish the AM workflow. 2. Identify data types and requirements. 3. Define data packages and identify where configuration changes occur. 4. Associate formats and establish a reference model for control.3.1. Step 1 - Establish the additive manufacturing workflowKim et al. [2] illustrate a typical AM workflow as presented in Figure 5. This workflow starts with design and planning activities, progresses to AM build planning and printing, then secondary manufacturing process and finally testing.274Figure 5: Illustrative additive manufacturing workflow [2].Through this process, the design as an entity changes from requirements and a conceptual idea to a digital model, then to machine instructions. These phases and the various stages in between are Digitally Orientated & Planning Focused.

These phases and the various stages in between are Digitally Orientated & Planning Focused. The design as an entity then transitions to a Physically Orientated & Production Focused phase whereby the design entity changes from the machine instruction into a physical green or as-built part and representative data of the physical production process. The design then changes through the various post-processing stages such as heat treatments and surface finishing processes and lastly testing and inspection of the final part. The physically orientated phase of the AM workflow contains an important data aspect as well. This data is the data generated and recorded during the physical production process, these data are required for verification and validation purposes.When formalizing the data flow, manufacturing follows a generic flow of design, process planning, manufacturing and quality evaluation.

Figure 6 presents the generic AM workflow that is decomposed from these four manufacturing functions as defined by Kim et al. [3]. This workflow serves as a reference workflow for all AM production runs and the basis for defining AM data requirements and packages.Figure 6: Reference additive manufacturing workflow functions.This AM workflow is performed at various phases of production, such as prototype, qualification and approved production as well as performed for multiple runs within these phases. This concept is presented in Figure 7. Data requirements can be defined for each of these blocks and data can be generated in each of these blocks. Data generated in blocks that follow in the lifecycle can utilize data from preceding blocks. In certain blocks not all functions in the AM workflow are performed, and example being during the prototype production phase, certain post-processing stages may not be required if only a design mock-up is needed.

When production transitions in either of the three dimensions, the data generated in the previous dimension shall be baselined and controlled. Reference to data blocks in this AM production framework can be made as defined in Figure 7. As an example, the heat treatment data from the second qualification production run would be referenced under the P2-A5.B data block.275Figure 7: Configurable “blocks” of data within the AM workflow.3.2. Step 2 - Identify the data types and requirementsThis step is an inclusive step, being that is not meant to satisfy a specific application scenario but instead establish allowable formats for meeting data requirements. The goal of this step is to map the “what” needs to be represented to available file formats on “how” this information can be represented. Figure 8 provides an example of the data that may be associated with a simplified workflow.

Figure 8 provides an example of the data that may be associated with a simplified workflow. At the highest level, the four-staged workflow in Figure 8 consists of Design, Preparation, Manufacture, and Inspection.Figure 8: Identification of the data types, the “WHAT”.276While the high-level progression of information is helpful to establish a general workflow and scope basic requirements, it is insufficient for establishing a detailed configuration management plan. Each stage of a workflow may consist of various configuration-control activities. Each of these activities may have unique data requirements depending both on AM-function and configuration-control types, and therefore unique data representations. Figure 9 provides an example of how data requirements are established based on the general workflow presented in Figure 8.

Figure 9 provides an example of how data requirements are established based on the general workflow presented in Figure 8. The sub-activities identified in Figure 9 are representative of activities that may drive data requirements, such as model tessellation and the generation of a build file. This step of the process is vital since it identifies the activities that may dictate data representation requirements.Figure 9: Deriving data requirements for the AM workflow.3.3. Step 3 - Define data packages and identify where configuration changes occurOnce the data requirements and types have been defined, this data is traced to distinct data packages at the various stages of the AM workflow. Therefore, moving from the WHAT to the HOW.

Therefore, moving from the WHAT to the HOW. The diagram in Figure 10 presents an example, demonstrating data packages for an AM workflow and the critical locations where data is transformed as well as where verification takes place.The data packages defined in this diagram are either transformed (changed) into other data packages or verified against previous data packages. This diagram defines the data packages that are the outputs of the AM workflow for a production phase, the configuration between design states and each of their production phases is additionally required. Data packages can be decomposed into sub-data packages or directly into their constituent data items.Identification of data inputs and outputs within the AM workflow is critical. Data is defined in either a digital or a mixed format such as documents. Physical entities are excluded from this diagram as their data representation is what is important here.

Data is defined in either a digital or a mixed format such as documents. Physical entities are excluded from this diagram as their data representation is what is important here. Referring to Figure 10, the “3D Design Model”, DP12, can be decomposed directly into data items whereas the “Post-Process Report” data package, DP51, can be decomposed into various sub-data packages such as Heat Treatment Report, Machining Report or Shot Peening Report and these further decomposed into their data items. Such data is application-specific, and the content varies greatly from one production instance to another.277Figure 10: Definition of data packages and changes along the AM workflow.The data packages and the decomposed data items identifiers lay the foundation for version control. When practicing configuration management, it is imperative to identify changes to the configuration.

When practicing configuration management, it is imperative to identify changes to the configuration. A change to a data item will require a change to that data package’s version identification to ensure integrity. Depending on the criticality of the data items or the data package itself, they can be classified as a configuration item and subsequently baselined and controlled. Figure 11 presents the decomposed data items of the DP11 and DP12 data packages.Figure 11: Exemplar definition of data items for the a) “Design & Customer Requirements”, and b) “3D Design Model” data packages.2783.4. Step 4 - Associate formats and establish a reference model for controlFigure 12 is an example of mapping “what” to “how”, where various design requirements from the design stages are met by various file formats. Additional details are needed to establish which file formats may best suit specific needs.

Additional details are needed to establish which file formats may best suit specific needs. Where step 3 was used to narrow down the data requirements for a given application scenario, step 4 will take this one step further by identifying allowable formats. Figure 12 illustrates how requirements can be mapped to formats at the most general levels, but an acceptable configuration management plan will require significant additional details on both the data requirements and the format capabilities.Figure 12: Basis of configuration management in digital production.The final step involves identifying applicable data formats for the defined data packages and items in Step 3, mapping the data requirements against the identified formats to ensure conformance and characterizing data integrity due to the transformations and verification between data packages and items.

Table 1 presents the description of a“Design & Customer Requirements” Data Package with identified applicable formats and Table 2 presents the same for the “3D Design Model” data package.Table 1: Exemplar description of the “Design & Customer Requirements” data package and applicable formats.DescriptionDesign and production requirements defined by industry and customer including standards and specifications.PurposeRequirements define what the customer wants and what industry requires produced. They are the first configuration baseline against which manufactured parts can be verified.FormatsReqIF, SysML, CSV, PDF, Text.Data ItemsData ItemAll AM?All Applications?Configuration Item?Rationale/CommentsDI111✔✔✔Each requirement needs an ID to reference it within the workflow and identify where and when it is verified.DI112✔✔✔The requirement description defines the details of what is required.

It is important that the description is unambiguous and that each person has the same understanding of what is required.DI113✔✔✔Various types of requirements can be defined. Requirements can apply to different stages of the workflow, be of different criticality and require different verification approaches.DI114✔✔✔Criteria for how the customer or industry requires the requirement to be verified or validated. These criteria are critical for production quality and release to the customer once produced.279Table 2: Exemplar description of the “3D Design Model” data package and applicable formats.Description3D design representation modelled in computer-aided software.PurposeThe 3D Design Model is the 3D presentation of what the customer wants manufactured.

It is used to represent certain requirements in a computer-readable format for analysis and further processing.FormatsSTEP, STL, IGES, OBJ, AMF, B-rep, 3MF.Data ItemsData ItemAll AM?All Applications?Configuration Item?Rationale/CommentsDI121✔✔✔The design ID is linked to the physical design body.DI122✔✔✔Design metadata can include data such as color, material, notes, reference to requirements etc.DI123✔✔The design state defines the stage of the design’s development such as prototype, qualification, or production.DI124✔✔✔The design body defines the data of the part’s solid geometry. This includes point cloud, boundary, and solid data.After each data package and their constituent data items are defined and applicable data formats are identified, the data formats are mapped against the data types and requirements in a matrix to identify conformance gaps and determine the conformant data formats as depicted in Figure 13.

When conformance gaps are identified, this indicates data from previous data formats need to be preserved.Figure 13: Map data to applicable data formats.Part of establishing a reference model for a specific scenario includes understanding the transitions between formats for the workflow scenario. As defined in Figure 10, there is a transformation and a verification event that occurs between the two example data packages, DP11 and DP 12, and therefore a configuration change has occurred. These are defined as T11-12 and V12-11 respectively. Table 3 describes the transformation between DP11 and DP12.280Table 1: T11-12 TransformationDescriptionTransformation of Design & Customer Requirements into CAD DesignPurposeCertain requirements are used to create the CAD design model. This transformation is largely performed by engineers and requires application-specific knowledge.Data IntegrityFrom requirements to a CAD design poses two key challenges.

This transformation is largely performed by engineers and requires application-specific knowledge.Data IntegrityFrom requirements to a CAD design poses two key challenges. The first being loss of information due to misunderstanding the requirements. The second challenge is that not all requirements can be fulfilled by a CAD design, this means certain requirements have to be stored along with the CAD design until they are addressed in the workflow.Conforming FormatsSTEP, AMF, 3MF.Nonconforming FormatsSTL, IGES, OBJ, BRep. These formats only partially represent the critical data items defined. If used, data will need to be preserved in additional formats.Table 4 presents the verification considerations for the exemplar case. The “3D Design Model” needs to be verified that it is conformant with the Design and Customer Requirements. Not all data packages are transformed and not all transformed data packages require verification.

Not all data packages are transformed and not all transformed data packages require verification. Data packages that have been verified are baselined and controlled.Table 4: V12-11 VerificationDescriptionVerification that the 3D Design Model data meets defined requirements.PurposeVerification that the design data meets the requirements is a critical step before continuing the workflow. This verification can be performed by various means such as simulations and design reviews. Once the design data is verified it is baselined and configuration controlled.Data IntegrityPotential integrity issues faced are the fidelity of the verification data generated and how well it represents the information it aims at verifying.

Additionally, not all information can be verified or to the same degree and certain data requires expert and application knowledge to verify.Establishing the transformations and understanding the verification needs is the final part of the configuration management plan. A user is left with application-independent scenarios, application-specific scenarios, identification of allowable formats, mappings of these formats to the scenarios, and finally an understanding of the impact of transitioning between formats throughout the workflow.DiscussionThe need for a standardized approach to the AM data workflow that addresses the full AM product lifecycle is identified as a current standardization gap [4]. The ISO/ASTM joint workgroup 73 was set up to address digital product definition and data management.

The ISO/ASTM joint workgroup 73 was set up to address digital product definition and data management. Supplementing the AMSC standardization gaps, the ASTM AM Center of Excellence published a strategic guide identifying the current challenges and gaps related to AM data management and schema as recognized within industry [9]. Noted in this guide is the need for principles defining what data is important to collect and how this data should be processed for utilization and understanding. Additionally, the need for a common data exchange format and determining minimum viable data packages for ensuring data pedigree and quality were identified as high priority topics.281The concept of a “data digital twin” as the data representation of the physical part and process is defined.

The MIL-HDBK-61B [8] point out that the concept of a digital twin is formed when these digital models evolve and mature until their behavior and performance become indistinguishable from their physical counterparts. This requires effectively defining configuration and stringent controls to realize this concept and its benefits.Data versioning and coding is an important aspect of configuration management. In modern computer-integrated and digital production processes, there is a substantial amount of data generated and recorded throughout the production lifecycle. This data is encapsulated in various formats, both digitally and physically, and controlling these various data configurations is ever more prevalent in AM production. The approach presenting in this paper demonstrates how this data can be packaged, assigned a code for referencing to the production phase, workflow location and version, and identification of applicable formats for representing this data.

This paper proposes an approach for implementing and addressing configuration management within the AM workflow. Implementing such configuration management principles is imperative for realizing and managing the digital twin and digital thread.ConclusionThe approach presented in this paper provides a framework for implementing configuration management practices in an AM organisation. The need for configuration management in digitalized production such as AM is justified and the implications this has on realizing the digital thread and twin are defined. A four-step approach is proposed consisting of 1) defining the AM production workflow, 2) identifying the data types and requirements, 3) defining data packages and identify the configuration changes between them, and 4) associate applicable data formats and determine conformant formats for data preservation.

nformal Cooling and Heating Channels using Laser Sintered ToolsNeil Hopkinson and Phill Dickens Rapid Manufacturing Research Group, De Montfort University, Leicester, LE1 9BH, UKwith thanks to the Warwick Manufacturing Group for sample Rapid Steel 2 parts and MH Tools for epoxy infiltration of DMLS bronze parts1AbstractThe EOS Direct Metal Laser Sintering (DMLS) and DTM Rapid Steel 2 processes may be used to create tools incorporating conformal channels behind the tool surface through which fluids may be passed. To date, a significant amount of work has been carried out to investigate the efficiency of using conformal channels to cool tools.This work suggests the use of conformal channels to both cool and heat a single tool. This may appear self-defeating at first but the selective nature by which conformal channels may make this a worthwhile means of generating hitherto unavailable thermal conditions within a tool.

This may appear self-defeating at first but the selective nature by which conformal channels may make this a worthwhile means of generating hitherto unavailable thermal conditions within a tool. Such conditions may then allow the successful production of geometries which had previously been impossible to mould.2Background2.1 The subject of conformal cooling channels has been a focal point for a large amount of research and published case studies in the field of Rapid Tooling. In terms of pure research, a number of academic institutions have published results from practical experiments and modelling work. Research at the University of Lowell (USA) investigated the positions of cooling channels in terms of their distance from the tool surface using stereolithography (SL) tools for injection moulding. The cross sectional shape and features designed to induce turbulent flow were also assessed.

The cross sectional shape and features designed to induce turbulent flow were also assessed. This research showed that the use of conformal cooling channels in SL tools could reduce temperatures by 170 deg F (77 deg C) (Janczyk et al 1997). Further research at De Montfort University (UK) has suggested that the low thermal conductivity renders conformal cooling channels redundant in SL tools (Colomer 1998 and Hopkinson 1999). Work performed at MIT showed that infiltrated steel tools produced with conformal cooling channels using the 3D Printing process resulted in cycle time reductions up to 15%. The same body of research showed that reductions in part shrinkage of 37% could be achieved with these tools using a standard cycle time (Sachs et al 1997). Further research from MIT investigated modelling the expected performance of conformal cooling channels and produced a systematic method to aid the design of cooling channels (Xu et al 1998).

Further research from MIT investigated modelling the expected performance of conformal cooling channels and produced a systematic method to aid the design of cooling channels (Xu et al 1998). The ExpressToolTM process was shown to reduce cycle times by 40% when using conformal cooling channels in electroformed tools (Macdonald 1998)Conformal Cooling Channels4902.2 Despite the widespread work on conformal cooling channels there appears to be little reported work regarding the use of conformal channels to heat a tool. Work at Leeds University (UK) has shown how conformal heating channels can be used to improve processing conditions in the moulding of elastomer compounds (Delgarno et al 2000).

Work at Leeds University (UK) has shown how conformal heating channels can be used to improve processing conditions in the moulding of elastomer compounds (Delgarno et al 2000). It seems logical that the use of conformal heating channels could be used successfully in other manufacturing processes which rely on thermal curing such as compression moulding and reaction injection moulding.Conformal Heating Channels2.3 The DMLS bronze process uses a laser to directly sinter metal particles layer by layer and the layer manufactured part is subsequently infiltrated with epoxy resin. This process has the significant advantage over processes such as Rapid Steel 2 in that no furnace operation for polymer burnout, sintering and infiltration is required. The DMLS process does however require higher temperatures during the laser sintering stage.

The DMLS process does however require higher temperatures during the laser sintering stage. A consequence of the differences between DMLS bronze and Rapid Steel 2 may be thought of as a “law of conservation of difficulty” issue; Rapid Steel 2 requires careful processing and thought in the post laser sintering stages whereas with DMLS bronze it is the laser sintering stage itself which requires the closest attention.Layer Manufacturing Technology used in this research3MethodologyThe aim of this programme of research is to investigate the possibility of using the DMLS bronze process to manufacture tools with heating and cooling channels and then to assess the performance of the tools. The first task is therefor to assess the types of geometries which can be built; having established the geometries which may be made it will then be possible to build tools and assess their performance.

For comparison, test parts were built using DTM’s Rapid Steel 2 process.3.1 Figure 1 shows the STL file for a test part to assess the manufacturability of conformal channels of different cross sectional shapes and sizes. The reason for building channels of different sizes was to assess the minimum size of feature which could be conformally cooled. The different cross sectional shapes were built as these would offer increased surface area for heat transfer (especially in the case of the star shaped channels) and easier manufacturabilty (in the case of the triangular shaped channels). Figure 2 shows the STL file for a test part to assess how easily loose powder could be removed from the layer manufactured model after laser sintering. In this part, longer channels with two right angled bends were used and again different cross sectional areas were included.Design of conformal heating and cooling channels49170mm40mmFigure 1.

STL file for part to assess different cross sectional shapes and areasFigure2. STL file for part to assess longer channels3.2 Figure 3 shows the design of the initial test moulding which was intended to assess how a tool with different wall thickness performs. The test tool was intended to assess firstly how the thinner wall section could be filled during injection and then to see how effective packing pressure could be applied – especially in the thick walled section which would be likely to show sink marks due to in-mould shrinkage. Finally, the test tool was intended to see how the variable wall thickness, with varying cooling rates, would be affected by warpage.Assess improvements in moulding quality4Initial Results4.1 Manufacture of Conformal Channels Figure 4 shows the conformal channel test parts built using DMLS bronze with different channels lengths, cross sectional areas and cross sectional shapes.

The two smallest diameter (2mm and 4mm) channels remained blocked with un-sintered powder after laser sintering. In the larger channels the loose powder was easily blown away with compressed air; these parts were not infiltrated with epoxy resin. The star shaped and triangular channels showed some crack propagation at sharp corners however these could probably be sealed with epoxy infiltration. The star shaped channels showed considerable delamination on downfacing surfaces suggesting the need for redesign, possibly using supports.49280sprue43621040All dimensions in mmFigure 3. Initial Test Moulding DesignCrack propagationDelaminationFigure 4. Conformal channel test parts built by DMLS after loose powder removal493Figure 5 shows the conformal channel test parts built using Rapid Steel 2. The test part with the long channels shows a hole where the green part broke while compressed air was used to clear loose powder in the 8mm diameter channel.

The test part with the long channels shows a hole where the green part broke while compressed air was used to clear loose powder in the 8mm diameter channel. Only the largest (10 mm diameter) hole remained unblocked. The different shaped cooling channels were built with no major difficulty although again, only the larger channels could be cleared of their loose powder.Hole blown in wall by compressed airFigure 5. Conformal channel test parts built by Rapid Steel 2 after loose powder removal.4.2 Manufacture of a test tool Figure 6 shows a sectioned picture of the STL file for the test tool core. The core includes two 10mm diameter conformal channels intended for cooling fluid around the 10mm thick wall and two conformal channels intended for heating fluid around the 2mm thick wall.Figure 6.

Cross sectional picture of the STL file for the test tool core494Figure 7 shows the core and cavity halves of the DMLS tool used to mould the test parts after it had come off the laser sintering machine and prior to infiltration with epoxy resin. This was the only tool used as a Rapid Steel 2 version was not builtFigure 7. The test tool core and cavity4.3 Moulding trials with the test tool After machining to allow for a sprue bush and fixing to a bolster, the DMLS tool was infiltrated with epoxy resin to seal its surfaces. The tool was then mounted on a 60 ton Battenfeld injection moulding machine and injected with polypropylene. Figure 8 shows a short shot with the thin section filling less readily that the thick section. The short shot, which had no follow up pressure applied was also subject to more warpage.Short shotFull shotFigure 8.

The short shot, which had no follow up pressure applied was also subject to more warpage.Short shotFull shotFigure 8. A short and full shot from the test tool495The conformal channels were not connected to heating or cooling fluids however thermocouples were located on two of the surfaces of the conformal channels. One thermocouple was used to measure the surface temperature inside the cooling channel next to the thick moulding wall section and one used next to the thin moulding wall section. These thermocouples were intended to monitor the amount of heat generated at the surfaces of the channels from the two different wall sections. Parts were injected at temperatures of 170 degrees C and 270 degrees C to assess how this affected the heat generated at the edges of the conformal channels.

Parts were injected at temperatures of 170 degrees C and 270 degrees C to assess how this affected the heat generated at the edges of the conformal channels. Figure 9 shows the temperature rises after injection of the melt in the two channels using the two different injection temperatures.37270oC injection near thick wall section35270oC injection near thin wall section33312927170oC injection near thin wall section170oC injection near thick wall section25020406080100Time after start of injection (seconds)Figure 9. Surface temperature readings from the inside of the conformal channelsFigure 9 shows that higher temperatures were recorded when injecting with a hotter melt as expected. Higher temperatures were recorded close to the thick wall section than the thin wall section as the thick section had more heat to dissipate.

Higher temperatures were recorded close to the thick wall section than the thin wall section as the thick section had more heat to dissipate. The actual rises in temperature however appear quite low when compared with previous work (Hopkinson, 1999); this may be due to the fact that the thermocouples were pressed against rather than embedded into the surface of the channels. Also, the effect of epoxy infiltration may result in the tools having a low thermal conductivity but a high specific heat. If the tools have a low thermal conductivity and high specific heat it may be difficult to achieve desired temperatures inside them however it should be relatively easy to maintain temperatures without too much fluctuation.4961205Conclusions and further workThe EOS DMLS bronze process may be used to build injection moulding tools with conformal heating/cooling channels with some restriction on their geometry.

In particular horizontal unsupported downfacing edges are susceptible to delamination and sharp corners can lead to crack propagation. Circular channels may be built and other shapes such as triangular and possibly fractals may be built to achieve different levels of heat transfer.The tools may be used for injection moulding however they must be sealed as they are porous after sintering. The slow thermal response measured in the heating/cooling channels suggests that a large amount of energy may be required to achieve desired temperatures throughout the tool; however, if achieved, these temperatures should be relatively easy to maintain. The next step to assess the thermal control will be to run tools with thermocouples embedded within the tool and then to run the tools with heating and cooling applied.

The next step to assess the thermal control will be to run tools with thermocouples embedded within the tool and then to run the tools with heating and cooling applied. The test moulding was produced relatively easily so further designs, perhaps with thinner sections, should be built to assess mould filling, the efficacy of packing pressure and warpage due to variable wall thickness.6ReferencesColomer, C., 1998, Business Benefits of Conformal Cooling, Final Year Dissertation for Bachelor’s Degree in Mechanical Engineering, De Montfort University,Hopkinson, N., 1999, Investigation into Part Ejection and Heat Transfer in the Direct AIMTM Process, Ph.D. Thesis submitted to De Montfort University, UK.Janczyk, R., McLaughlin,R.

nformal Lattice Structure Design and FabricationJason Nguyen1, Sang-In Park1, David W. Rosen1, Luis Folgar2, James Williams21School of Mechanical Engineering. Georgia Institute of Technology 2Paramount Industries, a 3D Systems CompanyAbstract One application of additive manufacturing is for fabrication of customized, light-weight material called Conformal Lattice Structures (CLS), a type of cellular structure with dimensions of 0.1 to 10 mm. In this paper, two advances are reported for designing CLS. First, computer-aided design technologies were developed for efficiently generating and representing CLS, given selected part model surfaces. Second, a method is presented for efficiently optimizing CLS by utilizing a heuristic that reduces the multivariate optimization problem to a problem of only two variables. The heuristic is: stress distributions are similar in CLS and in a solid body of the same shape.

The heuristic is: stress distributions are similar in CLS and in a solid body of the same shape. Software will be presented that embodies this process and is integrated into a commercial CAD system. In this paper, the method is applied to design strong, stiff, and light-weight Micro Air Vehicle (MAV) components.1Introduction1.1 Additive ManufacturingAdditive manufacturing (AM) refers to the use of layer-based additive processes tomanufacture finished parts by stacking layers of thin 2-D cross-sectional slices of materials. This process enable fabrication of parts with high geometric complexity, material grading, and customizability [1].Design for manufacturing (DFM) has typically meant that designers should tailor their designs to eliminate manufacturing difficulties and minimize costs. However, the improvement of AM technologies provides an opportunity to re-think DFM to take advantage of the unique capabilities of these technologies [2].

However, the improvement of AM technologies provides an opportunity to re-think DFM to take advantage of the unique capabilities of these technologies [2]. Several companies are now using AM technologies for production manufacturing. For example, Siemens, Phonak, Widex, and the other hearing aid manufacturers use selective laser sintering (SLS) and stereolithography (SLA) machines to produce hearing aid shells, Align Technology uses SLA to fabricate molds for producing clear braces (“aligners”), and Boeing and its suppliers use SLS to produce ducts and similar parts for F-18 fighter jets. In the first three cases, AM machines enable one-off, custom manufacturing of 10’s to 100’s of thousands of parts. In the last case, AM technology enables low volume production. In addition, AM can greatly simplify product assembly by allowing parts that are typically manufactured as multiple components to be fabricated as one piece.

In addition, AM can greatly simplify product assembly by allowing parts that are typically manufactured as multiple components to be fabricated as one piece. More generally, the technologies enable new opportunities for customization, unique capabilities of AM improvements in product performance, multi-functionality, and lower overall manufacturing costs. These unique capabilities include: shape complexity, where very complex shapes, lot sizes of one, customized geometries, and shape optimization are enabled; material complexity, where material can be processed one point, or one layer, at a time, enabling the manufacture of parts with complex material compositions and designed property gradients; and hierarchicalcomplexity, where multi-scale structures can be designed and fabricated from the microstructure through geometric mesostructure (sizes in the millimeter range) to the part-scale macrostructure.In this paper, we cover two main topics.

First, we present geometric construction methods that enable designers to take advantage of the shape complexity capabilities of AM processes. Specifically, we develop a method for constructing cellular materials that conform to the shapes of part surfaces; when restricted to lattice structures we call such constructs Conformal Lattice StructuresTM (CLS). The software that embodies this process is integrated into a commercial computer-aided design (CAD) system. Second, we present a design method, the augmented size matching and scaling (SMS) method, to optimize CLS efficiently and systematically.1.2 Cellular MaterialsThe concept of designed cellular materials is motivated by the desire to put material only where it is needed for a specific application. From a mechanical engineering viewpoint, a key advantage offered by cellular materials is high strength accompanied by a relatively low mass.

From a mechanical engineering viewpoint, a key advantage offered by cellular materials is high strength accompanied by a relatively low mass. These materials can provide good energy absorption characteristics and good thermal and acoustic insulation properties as well [3]. Cellular materials include foams, honeycombs, lattices, and similar constructions. When the characteristic lengths of the cells are in the range of 0.1 to 10 mm, we refer to these materials as mesostructured materials. Mesostructured materials that are not produced using stochastic processes (e.g. foaming) are called designed cellular materials. In this paper, we focus on designed lattice materials called meso-scale lattice structure (MSLS).In the past 15 years, the area of lattice materials has received considerable research attention due to their inherent advantages over foams in providing light, stiff, and strong materials [4].

Lattice structures tend to have geometry variations in three dimensions; some of our designs are shown in Figure 1. Deshpande et al. point out that the strength of foams scales as ρ1.5, whereas lattice structure strength scales as ρ, where ρ is the volumetric density of the material. As a result, lattices with a ρ = 0.1 are about 3 times stronger than a typical foam [5]. The strength differences lie in the nature of material deformation: the foam is governed by cell wall bending, while lattice elements stretch and compress.In order to effectively design cellular structures, we must be able to accurately model, determine the mechanical properties, and quantify the performance of these structures. Many methods have been developed to analyze various cellular structures. For instance, Ashby et al. has conducted extensive research in the area of metal foams [4].

Many methods have been developed to analyze various cellular structures. For instance, Ashby et al. has conducted extensive research in the area of metal foams [4]. Wang and McDowell have performed a comprehensive review of analytical modeling, mechanics, and various metal of characteristics honeycombs [6, 7]. Deshpande et al. have investigated extensively lattice the octet-truss cells, particularly analysis structure. However, assumed only the experience axial forces [5], while Johnson et al.

Deshpande et al. have investigated extensively lattice the octet-truss cells, particularly analysis structure. However, assumed only the experience axial forces [5], while Johnson et al. provided a more comprehensive analytical model of the truss structure by considering each0 0 00.5 0.5 0.51 1 1d d d1 1 1b) skin with single layer of lattice structure b) skin with single layer of lattice structurew w w0.5 0.5 0.5 0.5 0.5 0.5the strutsthat0 0 0 0 0 0v v v0 0 00.5 0.5 0.5u u u1 1 1a) octet-truss a) octet-trussc) Skin with 2 layers of truss structure made in SL c) Skin with 2 layers of truss structure made in SLFigure 1: Octet truss unit cell and example parts with octet truss mesostructure.strut as a beam experiencing axial, bending, shearing, and torsion effects.

The octet-truss structure was analyzed using a unit-truss model that consists of a node and set of half-struts connecting to the node [8].1.3 Design MethodsThe design synthesis method for cellular materials consists of size, shape, and topology optimization to address different aspects of the structural design problem. In order to understand optimization of structures, the definitions of three categories of structural optimization are important [9]. A typical size optimization involves finding the optimal cross-sectional area of each strut in a truss structure [10]. Shape optimization computes the optimal form that defined by the boundary curves or boundary surfaces of the body [11, 12]. The process may involve moving nodes to change the shape of the structure; however, the element-node connectivity remains intact. Topology optimization, according to Rozvany, finds optimal connective or spatial sequences of members or elements in a structure [13].

Topology optimization, according to Rozvany, finds optimal connective or spatial sequences of members or elements in a structure [13]. In topological optimization, the physical size, shape, and connectivity of the structure are not known. The only known properties are the volume of the structure, the loads, and the boundary conditions [9]. It can be seen that topology optimization involves aspects of both size and shape optimization. Three categories of structural optimization are illustrated in Figure 2. It can be seen that size and shape optimizations consider the material distribution in the structure to satisfy certain loading conditions while maintaining the same topology. On the other hand, the initial and optimal structures are completely different in the case of topology optimization. In this research, optimization variables of the truss structures are strut diameters. However, each unit cell of the MSLS can have a different configuration depending on the selection criteria.

In this research, optimization variables of the truss structures are strut diameters. However, each unit cell of the MSLS can have a different configuration depending on the selection criteria. Therefore, “topology optimization’ will be the term used in this research for designing and optimizing MSLS.for cellular structures dates as far back as a century ago. In 1904, George Michell, an Australian engineer, published a theory that defines the existence of an analytically optimal truss structure under certain loading conditions [14]. However, Michell trusses two dimensions and are not conducive to conventional manufacturing due to varying lengths and curved beams needed for optimal solution.

However, Michell trusses two dimensions and are not conducive to conventional manufacturing due to varying lengths and curved beams needed for optimal solution. Hence, it is very limited in application.StructuraloptimizationFigure 1: (a) Sizing optimization of a truss structure, (b) shape optimization, and (c) topology optimization [9]arelimitedtoThe topology optimization techniques used to design truss structures are based on one of two approaches: the homogenization (continuum) approach and the ground (discrete) truss approach. The homogenization approach in topology optimization is a material distribution method that considers the design space as an artificial composite material with an infinite number of periodically distributed small holes. The problem is transformed from a topology optimization problem to a sizing optimization problem by considering the sizes of these small holes as design variables.

The problem is transformed from a topology optimization problem to a sizing optimization problem by considering the sizes of these small holes as design variables. The main task is to create a microstructure model using a material density function. In the final optimal structure, regions with density at or near one are filledwhile regions with density at or near zero are empty. The method was pioneered by Bendsøe and Kikuchi in 1988 [15].The ground truss approach starts with a ground structure, which is a grid of all elements connecting the nodes in the design space. The optimal truss structure is realized by selecting an optimal substructure from this pre-defined ground structure. Ultimately, the ground-truss approach is a sizing optimization problem, where the cross-sections of ground truss members are the continuous design variables for the optimization. The cross-sections of the struts are sized to support the applied loads on the structure.

The cross-sections of the struts are sized to support the applied loads on the structure. Struts with cross-sections near zero are then removed to obtain the optimal structure [16].2 CLS Design MethodThe basic idea of how cellular materials are created is presented here. Four example primitive cell types are shown in Figure 3, three of which are lattice structures and the fourth is a foam. These cell types are 2-dimensional for simplicity of presentation. The octet lattice in Figurea is an example 3-D cell type. Lattice structures consist of a set of struts (beams) that connect the nodes of the lattice.a) uniform lattice a) uniform lattice a) uniform latticeFigure 3: Cellular primitives: three lattice structures and one web structureb) conformal lattice b) conformal lattice b) conformal latticeFigure 4: Uniform and conformal lattice structuresTo generate the cellular designs in Figure, the primitive cell types must be mapped into a mesh.

In 2-D, the mesh consists of a set of connected quadrilaterals. In 3-D the mesh consists of hexahedra (6-sided volume elements with planar sides). The uniqueness of our work is our use of conformal cellular structures, rather than uniform “lattice block” materials, that can be used to stiffen or strengthen a complex, curved surface. To see the difference between conformal and uniform structures, Figure 4a is an example uniform lattice structure, while Figure 4b shows a conformal lattice. Meshes for uniform structures consist of cube elements in 3-D (squares in 2-D), while for conformal structures, the mesh elements are general hexahedra. We have developed a new algorithm for generating conformal meshes that are used to create conformal lattice and cellular structures. An older algorithm based on a mapped meshing approach [18] has been updated significantly. We prefer that mesh elements are as cubic as possible; i.e., are of uniform thickness and uniform size.

An older algorithm based on a mapped meshing approach [18] has been updated significantly. We prefer that mesh elements are as cubic as possible; i.e., are of uniform thickness and uniform size. Such meshes are typically not generated by the free meshing methods in finite-element analysis codes, while typical part geometries are too complex for mapped meshing methods.2.1 CLS Construction MethodThe overall method for generating conformal cellular structures is shown in Figure 5. It consists of two main steps, indicated by the shaded rectangles: computing 3D conformal mesh, and populating the mesh with cells. Inputs and outputs of the steps are shown as ovals. Each step is detailed below.The objective of the meshing algorithm is to generate a conformal hexahedral mesh into which cells from the cell library can be placed. One or more layers of cellular structure can be placed to support the part’s skin.

One or more layers of cellular structure can be placed to support the part’s skin. The input to the algorithm may be a CAD solid model of the part, a surface model of the part, or a triangulated surface model (STL file) of the part. A method of constructing solid and STL models of lattice structures was presented in [17]. The method utilized the conformal lattice generation algorithm from [18].18 1818 1816 1616 1614 1414 1412 12Z Z12 1210 1010 108 88 86 66 60 04 42 20 00 02 24 4X X2 24 45 50 0Figure 5: CLS construction method2.2 Construct 3D Conformal MeshThe algorithm to generate a 3-D conformal mesh is shown in Figure 6. The first step is to divide the part boundary into relatively flat regions, since it is easier to control the mesh generation method if regions are not very curved.

The first step is to divide the part boundary into relatively flat regions, since it is easier to control the mesh generation method if regions are not very curved. We implement an absolute angular deviation measure between surface or triangle normal vectors to determine if that surface or triangle should be added to the region being generated. For the purposes of presentation, we will assume that a STL file was given. Then, one triangle is chosen as the first triangle of a region. The normal vector of each connected triangles is compared to the normal of the first triangle; if they differ by less than a given tolerance, the triangle is added to the region. As an example, the tolerance for the simple part in Figure 5 was chosen so that the model consists of two regions: the cylindrical surface and the planar surface.For each region, three main steps are performed as indicated by lines 3, 4, 5 of Figure 6. The first is to compute the offset of the object boundary.

The first is to compute the offset of the object boundary. An offset is a collection of points that are at a specified distance away from the starting surface (distance is called the offset distance). Asan example, an offset of a circle is a circle that is concentric with the first circle. Generally speaking, a positive offset results in a larger object, while a negative offset results in a smaller object. We use an offset method developed for tessellated part surfaces [19], but any offsetting method could be applied.Algorithm Construct Conformal Mesh Input: CAD or STL model, desired element size Output: hexahedral mesh that conforms to the outer surface of the given model 1. Partition part model into relatively flat regions. 2. For each region, 3. compute offset of region boundary 4. construct tri-parameter volume 5. divide parameterized volume into hexahedra 6. End for each region. 7.

2. For each region, 3. compute offset of region boundary 4. construct tri-parameter volume 5. divide parameterized volume into hexahedra 6. End for each region. 7. Ensure region boundaries match.Figure 6: Algorithm for constructing a conformal meshThe second step is to construct a tri-parameter volume between the original surface and its offset. Surfaces are parameterized using two parameters. Volumes are parameterized using three parameters. This step of the algorithm imposes a surface parameterization on the region of the part that will be reinforced [20]. The same parameterization is transferred to the offset surface. Finally, a lofted volume is constructed from the original surface to the offset surface.The third step is to generate the conformal mesh by dividing the parameterized volume into individual elements (hexahedra). This is very straightforward.

This is very straightforward. After selecting either the number of elements or their typical size, increments in each of the three parameters are computed, which are used to generate elements of the mesh. For example, if the increment in u is chosen to be 0.1, then 10 elements in the u direction will be generated, since (1 - 0) / 0.1 = 10. By successively incrementing each of the three parameters that define the tri-parametric volume, mesh elements are created.The final step in Algorithm Construct Conformal Mesh (line 7) is to ensure that region boundaries match by adjusting node positions and by adding elements, if necessary. Since the regions are parameterized separately, the hexahedral elements may not match well. Nodes from neighboring regions may be moved and merged to achieve matched boundaries.

Since the regions are parameterized separately, the hexahedral elements may not match well. Nodes from neighboring regions may be moved and merged to achieve matched boundaries. Also, a series of hexahedral or tetrahedral elements may be added in gaps between meshes in neighboring regions.2.3 Construct MSLSThe algorithm for the second step (populate mesh with cells) of the overall conformal cellular structure design method from Figure 6 is described here. One input to the algorithm is the conformal hexahedral mesh that was generated in the first step. The other input is the cell types contained within a library. The first step is to partition the mesh elements into regions such that within each region the loading conditions are similar on each element. These need not be the same regions that were used for mesh construction. For each region, a cell type from the cell type library is selected to populate the mesh elements in that region.

These need not be the same regions that were used for mesh construction. For each region, a cell type from the cell type library is selected to populate the mesh elements in that region. The idea is to match the region’s loading conditions to the cell type, such that the cell type is effective at supportingthe loading conditions. In this manner, the resulting cellular structure is more likely to be lighter for a given level of stress or deflection. These operations may be performed concurrently or maybe performed sequentially, depending upon the designer’s preference and may be automated or be performed by the designer directly.The final step in Figure 7, “Apply Selected Cell Types to Selected Mesh Elements,” is where the actual cellular geometric model is constructed. This operation has been called population of mesh elements earlier. One simply maps a cell type into a mesh element.

This operation has been called population of mesh elements earlier. One simply maps a cell type into a mesh element. Since both the cell type and the mesh element are defined parametrically, a simple parametric mapping algorithm can be applied to directly construct cell geometry.To construct a STL of the CLS, additional geometric construction operations must be performed. We utilize the approach described in [17], where solid models of the half-struts incident at each node of the mesh are constructed using Boolean operations in a solid modeling system. Then, each solid “node” is tessellated and the triangles are written to a STL file.The resulting conformal cellular geometric model can be subjected to optimization methods in order to reduce weight, increase strength, increase stiffness, achieve some other objective, or achieve some combination of objectives.

Such methods have been applied by a number of research groups [21].2.4Integration into CADSet general parametersThe overall process for constructing MSLS has been embodied in a software package called TrussCreator, which is available as a plug-in for the Siemens NX CAD system. The construction procedure is shown in Figure 7. For each step, dialog boxes are designed and created for user input. The user can access these dialog boxes through the TrussCreator menu, as shown in Figure 8. In the lattice parameter setting menu, input information about diameter of struts, the size of unit cells, and partition angle (step 1 in Figure 6) for given surfaces, and other settings.the user canNoSet lattice parametersSelect lattice typeSelect surfacesSet lattice directionLast part?YesGenerate lattice - TrussCreatorDisplay latticeAfter entering the parameter settings, the user can select the lattice structure type from a set of pre-defined types.

In the surface selection stage, the class selection dialog is loaded that makes selection for only surfaces in the current working part. After selecting the surface, the user can choose what direction to add the lattice structure. Arrows appear that indicate the normal direction in which the lattice structure will be added; the direction can be flipped by selecting the surface that user wants to change. TrussCreator generates lattice structures based on input information mentioned above. After generating the lattice structure, it is displayed with its parent surfaces in the NX system. The user can examine the lattice structure details using NX functionality and save the structure as an STL file. Capabilities are offered to edit the lattice structure (add/modify/delete nodes and edges). Additionally, structural analysis or optimization can be performed by exporting files for input into ANSYS.Save STL fileFigure 7: Interface flow chartFigure 8.

Additionally, structural analysis or optimization can be performed by exporting files for input into ANSYS.Save STL fileFigure 7: Interface flow chartFigure 8. TrussCreator menu.3 Augmented Size Matching and Scaling (SMS) MethodRegardless of which structural optimization approach outlined in Section 1.4 is used for the design of meso-scale lattice structures (MSLS), an actual multi-variable optimization routine must be performed. Since the computational complexity of the design problem often scales exponentially with the number of design variables, topology optimization is infeasible or impractical for large design problems. The Size Matching and Scaling (SMS) method uses a heuristic to reduce the multivariable optimization problem to a problem of only two variables [22]. The heuristic is based on the observation that the stress distribution in a MSLS will be similar to the stress distribution in a solid body of the same overall shape.

The heuristic is based on the observation that the stress distribution in a MSLS will be similar to the stress distribution in a solid body of the same overall shape. Based on the computed local stress states from the solid body analysis, unit cells from a predefined unit-cell library are selected and sized to support those stress states. The optimal diameters of these struts are then computed by performing a two-variable optimization. This design approach removes the need for a rigorous multi-variable topology optimization, which is a main bottleneck in designing MSLS. The previous SMS method [22] was limited to MSLS designs with simple geometry and shape because of the lack of systematic ground structure generation capability.

The previous SMS method [22] was limited to MSLS designs with simple geometry and shape because of the lack of systematic ground structure generation capability. To overcome this limitation, the new augmented SMS method presented here integrates the CLS construction methods outlined in Section 2 to the ground structure generation process of the SMS method.3.1 Problem FormulationEach meso-scale lattice structure design problem has its own loading condition, geometric properties and desired performance specification. However, they can all be characterized as multi-objective design problems using the Compromise Decision Support Problem (cDSP) method [23]. The specific problem formulation for the SMS method is presented in Table 1.The symbols pBG, pF, pM represent the boundary, loading and material properties, respectively. The strut diameter, Di, can either range from the lower diameter bound, DLB, to the upper diameter bound, DUB, or zero.

The strut diameter, Di, can either range from the lower diameter bound, DLB, to the upper diameter bound, DUB, or zero. The symbol σi represents the axial stress value in each i strut. The symbols V and d represent the volume and the deformation of the structure. Wd and WV represent weighting variables for d and V in the minimization function, Z. The volume of the structure is calculated by summing the volume of all the struts in the structure, which are assumed to be cylinders:V∑ π=×2 D i 4×l i(1)where Di and li represent the diameter and length of each of the i strut in the structure. In this calculation, the overlapping volumes where the struts meet are not subtracted from the overall volume of the structure because they are assumed to have negligible contributions in order to simplify the calculation.

External sources of information include the unit-cell library and the solid- body finite element analysis. Using that information, the determination of the strut diameters, shown in (a) of Table 1, reduces to a 2-variable optimization problem. It can be seen that Di,k can be determined using the pre-scaled maximum and minimum diameter value, DMAX and DMI(cid:28), i jS , and a unit-cell scaling factor from the solid- a stress scaling factor from the unit-cell library, i jS . Hence, only DMAX and DMI(cid:28) need to be determined through body stress analysis, optimization. The minimization function, (g) of Table 1, is formulated in the least-squares format to minimize the deflection of the structure, d, and deviation of the structural volume from a target volume, Vt.

The minimization function, (g) of Table 1, is formulated in the least-squares format to minimize the deflection of the structure, d, and deviation of the structural volume from a target volume, Vt. Wd and WV represent the weighting variables for d and V.u ,L ,The optimization process of DMAX and DMI(cid:28) requires calculation of deflection, volume, and associate stresses using finite element analysis of the truss structure. The finite-element package, which assumes each truss member as a beam element, was developed in MATLAB [24]. Once the optimization is done, the diameter of each strut is obtained using Equation (a) of Table 1. The optimized maximum and minimum diameters of the structure are denoted as Dmax and Dmin to differentiate from the pre-scaled maximum and minimum diameter value, DMAX and DMI(cid:28). It is important to note that the finite element analysis of the truss structure is conducted using the scaled/true diameters of the structure.

It is important to note that the finite element analysis of the truss structure is conducted using the scaled/true diameters of the structure. Other problem formulations with different objective functions can also be used with the SMS method.3.2 Overview of Augmented SMS methodThe SMS method can be divided into eight discrete tasks in seven that are completed steps, as summarized in Figure 9. Outputs of each step are shown in the shaded box under each step.2a. Generate Ground StructureUnit-Cell Model1. Specify geometric, boundary and material propertiesDesign Parameters3. Correlate FEA nodes to Unit-cell Model2b. Generate and Analyze Solid ModelNodal Stress Results3.2.1 Step 1: Specification of loading, boundary conditions and material propertiesMapped Nodes4. Normalize Stress ResultsUnit-Cell Stress ValueIn this first step of the method, boundary the conditions, material properties, are and specified for the target meso- scale lattice structure.

Normalize Stress ResultsUnit-Cell Stress ValueIn this first step of the method, boundary the conditions, material properties, are and specified for the target meso- scale lattice structure. These properties will be utilized to perform the stress analysis of both solid-body representation in step 2b and the truss structure during the optimization process of step 7. These values include the material properties such as Poisson’s ratio, Young’s modulus of elasticity, and the desired loading and boundary conditions.5. Generate Unit-Cell TopologyUnit-Cell LibraryStructural Topologyloadingconditions6. Remove Unessential/ Duplicated StrutsClean Topology7. Determine Diameter ValueOptimized StructuretheFigure 9: Overview of augmented SMS method3.2.2 Step 2a: Generation of ground structureIn this step of the method, the ground structure of the MSLS is created.

The ground structure only specifies the bounding geometry of the truss structure and contains no actual struts or materials. In this implementation of the SMS method, a free mesh approach is utilized to generate the ground structure that conforms to an arbitrary complex surface. Computer-aided technologies were developed for efficiently generating and representing the lattice structure [25], as described in Sec. 2.2.3.2.3 Step 2b: Solid body finite-element analysisA solid body is generated that envelopes the part model surfaces and the MSLS and a stress analysis is performed using finite-element analysis. The loading and boundary condition, and material properties for the structural analysis are specified in step 1 of the method. The purpose of this step is to obtain the stress distribution of the solid-body structure and extrapolate this information to determine the stress distribution and the local stress states in the MSLS.

Once the analysis is complete, the von Mises stress distribution of the structure is obtained. The primary deliverable of this step is the general state of stress at each node, which is characterized by six independent normal and shear stress components, σxx, σyy σzz, τxy, τxz, and τyz.3.2.4 Step 3: Map FEA nodes to ground structureIn order to use the finite-element analysis result obtained from step 2b, the stress results must be appropriately mapped to the ground structure. The goal of this step is to determine which finite-element nodes correlate to which unit-cell region in the ground structure. Sincethe free mesh approach is utilized to generate the ground structure, the augmented SMS method requires a FE node classification method that works with arbitrarily shaped mesh elements.The process starts by dividing each face of the unit cell into triangles, since triangles are convex and planar, and every polygon can be broken up into a set of triangles.

An outward- pointing normal for each triangle can then be obtained by computing the cross-product of two of the edges. After all the normal vectors are obtained for the unit cells, we can determine whether or not a finite-element node falls into the unit cell by computing the dot-products between the triangle normal vectors and the vector from a vertex of each triangle to the node. For hexahedra mesh elements, there will be a total of twelve triangles with twelve outward- pointing normals and twelve dot product operations. If and only if all the dot product results are either 0 or less than 0, then the finite-element node belongs to that unit cell. In the case that one of the dot products is equal to 0, the finite-element node is on the boundary between multiple unit cells and will be included in each.

In the case that one of the dot products is equal to 0, the finite-element node is on the boundary between multiple unit cells and will be included in each. After the node mapping process is done, each unit cell will contain a list of finite element nodes that will be included in the calculation of the stress distribution in that unit cell.3.2.5 Step 4: Stress Scaling and (cid:28)ormalizationAfter step 3 is complete, the stress values from the finite-element nodes in each unit cell are averaged to determine average stress values of six independent normal and shear stress components. Only the absolute values of the stresses are averaged. The stress results from FEA are only relevant for the solid-body structure. Instead, the stress distribution is of interest and will be used to guide the setting of strut sizes. Therefore, the stresses are normalized from zero to one such that the largest value of stress is equal to one.

Therefore, the stresses are normalized from zero to one such that the largest value of stress is equal to one. These six scaling values correlate to six entries of each configuration in the unit cell and will be utilized to size the struts during the topology generation process in step 5.In topology generation, the diameter values of the selected unit-cell configuration from the preconfigured unit-cell library are scaled against the associated stress values (σxx, σyy, σzz, τxy, τxz, and τyz) and then mapped to the unit cells in the ground structure. However, since the solid-body results are provided relative to the global coordinate system, stress transformations are needed to ensure correct topology generation, which is performed by rigid-body rotation of the axes. However, since the unit cell is not necessarily a cuboid hexahedron, a representative local coordinate system can be determined using the following approach.

However, since the unit cell is not necessarily a cuboid hexahedron, a representative local coordinate system can be determined using the following approach. Each unit cell, such as in Figure 10, from the ground structure is characterized by 8 nodes and 12 edges. Three edges of the unit cell, edge 1-2, edge 1-4, and edge 1-5, respectively, are selected as reference edges, nominally representing the x, y, and z axes, respectively. For each edge that corresponds to the direction of a certain axis, the angle between that edge and the corresponding reference edge is calculated; e.g. in the x-axis direction, the angles between edge 1-2 and edge 5-6, edge 1-2 and edge 8-7, and edge 1-2 and edge 4-3 are calculated and averaged. This step is repeated for the other two directions. The reference edge with the lowest average angle is selected as the starting axis for that particular direction. The reference edge with the second lowest average angle is selected as a second axis.

The reference edge with the lowest average angle is selected as the starting axis for that particular direction. The reference edge with the second lowest average angle is selected as a second axis. The cross product is performed between the first and second axes to find the third orthonormal axis. Then another cross product is computed between the third and first axes gives the second orthonormal axis. This approach determines the local orthonormal coordinate system of a unit cell from the ground structure.Figure 10: Unit-cell regionFigure 11: Rotation of coordinate system (x’, y’, z’)After obtaining the local coordinate system for the unit cell, the relative orientation between the local and global coordinate system can be determined.

The global coordinate system, xyz, and local coordinate system, x’y’z’, are shown in Figure 11 where α1 is the angle between the x’ and x axes, β1 is the angle between x’ and y axes, is the angle between x’ and z axes, α2 (not shown) is the angle between the y’ and x axes, and so forth.Let R be the rotation matrix that transforms the vector components in the original coordinatesystem to those in the primed system, then' x     ' y     ' z  =    R 11R21R31R 12R22R32R 13R23R33x       y       z   (2)From Eqn. 2, it can be seen that the unit vector x’ can be expressed in the original coordinate system as'x=R x R y R z 12 13++11(3)where Rij are direction cosines. Therefore, x’ can be expressed in terms of x, y, z using Eqn. 4.x'=cos( α 1)x+cos(β 1)y+cos( θ 1)z(4)Similarly, axes y’ and z’ can be expressed in term of x, y, and z.

This stress tensor is defined in the deformed state of the material and is known as the Cauchy stress tensor [26]. With the rotation matrix given in Eqn. 5, the Cauchy stress tensor in the local coordinate system (x’, y’, z’) can be obtained using Eqn. 7.[ σ] '[ ] σ= RRT(7)RT is the transpose of R, σ is the Cauchy stress tensor in global coordinate system (x ,y, z), and σ' is the Cauchy stress tensor in the local coordinate system (x’, y’, z’). This follows the rule of changing second-order tensor components under rotation of axes [26]. Ultimately, there will be six stress values for each unit cell, σxx, σyy σzz, τxy, τxz, and τyz, which correspond to the scaling factors,, in (a) of Table 1.3.2.5 Step 5: Topology generationThe unit-cell lattice structure selection and mapping process will be described in detail in Section 3.3. After this step is complete, the structure will have a topology designed for the anticipated stress distribution in the truss structure.

After this step is complete, the structure will have a topology designed for the anticipated stress distribution in the truss structure. The relative thickness of one strut to another is known. However, these normalized diameters must be correlated with actual strut diameter values in step 7 of the method.Figure 12: General state of stress [26]3.2.6 Step 6: Ambiguity resolutionSince the unit cells are populated individually, therewill be instances of overlapping struts between adjacent unit cells. These struts will have identical start and end nodes. To resolve this ambiguity, the largest diameter strut is kept and all other smaller struts are removed. Duplicated nodes are also removed.3.2.7 Step 7: Diameter SizingThe normalized strut diameters must be replaced with the actual diameter values to satisfy the loading and volume condition.

It is rewritten in Eqn. 9 as a function of both DMI(cid:28) and DMAX, where V(DMI(cid:28), DMAX), volume, and d(DMI(cid:28), DMAX), deformation, are functions of only DMI(cid:28) and DMAX. Deformation, d, represents anyunit of measure that is directly proportional to structural stiffness, such as tip deflection or strain energy. The target structure must attempt to minimize both volume and deflection. However, these two goals have competing effects. The target deflection is usually set to zero. Two algorithms used to perform this two-variable minimization are the Levenburg-Marquardt and active-set algorithms.

The target deflection is usually set to zero. Two algorithms used to perform this two-variable minimization are the Levenburg-Marquardt and active-set algorithms. The Levenburg-Marquardt algorithm has documented success in design and optimization of MSLS [21], while the active-set algorithm is documented to have success in optimization of multivariable, nonlinear and constrained optimization problems [29].( Z DMI(cid:28),DMAX)=( × W d D(dMI(cid:28),DMAX))2+  W V×( V DMI(cid:28), D V tMAX)−V t  2(9)In addition to exploring the two-variable optimization approach, it was noted in previous research that for a particular truss structure there is an ideal relationship between DMI(cid:28) and DMAX such that when the ratio is approximately equal to 28% for a specific target volume, the structure would have the least deflection [27].

This finding has significant effect because it would reduce the two-variable equation involving DMI(cid:28) and DMAX to a one-variable equation. As a result, the equation to determine the diameter of each strut becomes:D, i k= 0.72(Su , i j×SL , ij)+0.28 ×DMAX(10)In this research, both one-variable and two-variable approaches will be used. The results willbe compared in terms of the deformation and design time.3.3 Unit-cell LibraryThe second component of the augmented SMS method is the unit-cell library, which was developed to generate the topology for the MSLS [22]. There are seven different unit-cell configurations in the library. Each configuration has six entries with each specialized for six independent normal and shear stress components.

There are seven different unit-cell configurations in the library. Each configuration has six entries with each specialized for six independent normal and shear stress components. Entries in the library were optimized for loading conditions corresponding to each of the six stress states.3.3.1 The optimization processThe problem formulation utilized for the optimization of unit cells is shown in Table 2.Table 2: Qualitative cDSP formulation for the optimization of unit cells [15]Given: Find:Loading and Fixity Conditions, Starting Lattice Topology Truss Diameters/Lattice TopologySatisfy:Target Strain Energy Maximum Stress ValueMinimize: VolumeFor unit-cell optimization, the objective is to minimize the volume of the unit cell. The stiffness is set as a constraint to force the performance of all the optimized unit-cells to be equal.

The stiffness is set as a constraint to force the performance of all the optimized unit-cells to be equal. Strain energy ∆U, is the metric to measure stiffness, which is calculated as:△U=△F 2d(11)whereF△ is the average magnitude of the load and d is the total displacement of the structure. 2Strain energy is widely used in topological optimization problem.The optimization process for the unit cell is divided in to five separate steps as describedbelow.Step 1: Insert Initial Unit-Cell Configuration: Each unit cell is defined in a cuboid region by 8 nodes in each corner. A diameter value of “1” is assigned to all the struts in the unit cell.Step 2: Apply Loading Conditions: In this step, the unit cell is loaded with six loading conditions, each for a component of the stress state. The loading conditions are shown in Figure 13.These loading conditions must be applied in multiple directions.

The loading conditions are shown in Figure 13.These loading conditions must be applied in multiple directions. For instance, there are 4 shear directions in the XY plane including τxy, -τxy, τyx, -τyx. The unit-cell must be optimized individually then combined to form the final optimized unit-cell.for each direction andStep 3: Optimize unit-cell: After defining the base-lattice structure and loading conditions, the unit-cell was optimized using the parameters in Table 3. The analysis and optimization was done in ANSYS 13. 0Figure 13: Loading conditions for unit-cell optimization [22]Step 4: Combine Optimized Unit-Cells: Since the loading conditions are applied in multiple directions, the results from each direction must be combined to form an optimized unit cell. The process is illustrated in Figure 14 for shear stress in the XY plane.

The process is illustrated in Figure 14 for shear stress in the XY plane. When combining the results, the largest diameter for each strut is kept and all other instances are deleted.After all the configurations are optimized, the diameters of the unit cells are normalizedfrom 0 to 1:Dnorm j k l , ,=Dj k l , , max D l(12)Table 3: Optimization parameter for unit-cell optimization in ANSYS[22] 50Strain Energy Constraints (mJ)Poisson Ratio0.3where j represents each strut for each kth configurations l stress for each of directions. k goes from one to seven because there are seven configuration. l goes from one to six because there are six stress directions, σxx, σyy, σzz, τxy, τxz, and τyz. Hence,Elastic Modulus (N/mm2)1960theLoading Magnitude (N)10Element TypeBEAM 4lD becomes the unit-maxrepresents the largest diameter value among the six stress directions.

Hence,Elastic Modulus (N/mm2)1960theLoading Magnitude (N)10Element TypeBEAM 4lD becomes the unit-maxrepresents the largest diameter value among the six stress directions. cell library scaling factor that is used in the diameter determination step.After the normalization process is complete, the unit-cell library is stored in a list with threekey parameters: the nodal coordinates, the elements, and the diameters of each element3.3.2 Unit-cell selectionOne way to generate the best topology for the structure is to iteratively populate each unit cell in the ground structure with a configuration from the library and analyze the performance of the structure. However, it is computationally infeasible because there are MN number of possible combinations of topology where M is the number of configurations in the unit-cell library and (cid:28) is the number of unit cells in the ground structure.

For instance, a ground structure with 5 unitcells would already have 75 = 16807 unique topologies. Therefore, a heuristic was developed for the selection process.Since all the configurations of the unit- cell library are optimized such that they have identical performance, the structure with the smallest normalized volume is is performed selected. using the Equation (18) [22], which incorporates three terms. The first is the lattice structure unit-cell sum of all volumes from the 6 stress states. The second term is the volume of the structure if it were mapped into the particular mesh element, while the third term represents a performance term based from empirical results.

The second term is the volume of the structure if it were mapped into the particular mesh element, while the third term represents a performance term based from empirical results. This term attempts to predict how well a configuration will perform when multiple instances of the configuration are utilized.The selectionFigure 14: Combination of optimized unit-cells for shear stress in XY plane for Cantley configuration [22]This selection process is performed for each unit cell from the ground structure and theconfiguration with the lowest rating, r, is selected for that unit cell.

∑ P∑(13)= r W v×× V W V W++×σvnnetpwhere∑σ = VV xx+Vyy+V zz+V xy+V xz+Vyz(14)∑= P P xx+P yy+P zz+P xy+P xz+P yz(15)and Wv, Wvn, and Wp are weights (importances) on each term.For each topology configuration, the volume of each entry in six primary stress directions (σxx, σyy, σzz, τxy, τxz, and τyz) is multiplied by the corresponding normalized stress results from step 4 of the SMS method and then summed to determine Vσ∑ as shown in Eqn. 14. For each topology configuration, Vnet each configuration calculated by combining all six entries in six stress direction and removing the overlapping struts. P∑ for each topology configuration is determined, as shown in Eqn. 15, using a performance table. These values are provided in Table. They are determined using results from a design example [22].

P∑ for each topology configuration is determined, as shown in Eqn. 15, using a performance table. These values are provided in Table. They are determined using results from a design example [22]. In this example, a 15 × 15 × 15 cm cube is divided into 3 × 3 × 3isthe net volume ofFigure 15: Unit-cell library [22]of the same unit-cell configuration. The same loading and boundary condition, shown in Figure 13, is applied to the cube. There are six loading condition approximating the six axial and shear stresses. Each of these loading conditions will be applied to all seven configurations in the library; therefore, there will be 42 unique topologies, as shown in Figure 15. The strain energy from each topology is calculated. The results are normalized between 0 and 1 and plugged in the performance table to calculate P∑.

The strain energy from each topology is calculated. The results are normalized between 0 and 1 and plugged in the performance table to calculate P∑ . In addition, there are three weighting values Wv, Wvn, and Wp, which can be set manually to vary the importance of each contributing factors to generate different topologies.

In addition, there are three weighting values Wv, Wvn, and Wp, which can be set manually to vary the importance of each contributing factors to generate different topologies. The configuration with the lowest rating is selected and mapped to that particular unit cell in the ground structure.Table 4: Performance table used for selection of unit-cell configuration [22]XX Axial YY Axial ZZ Axial XY Axial YZ Axial XZ AxialCrossed0.07450.06930.03750.08100.07470.0752Cantley0.53990.48850.05390.54180.53530.2626Octet0.22810.20230.10500.10040.08910.0863Paramount10.01970.09070.05000.98650.39040.3734Diagonal0.07430.07040.03900.11660.08810.0956Paramount21.00001.00001.00000.60430.55690.5462Midpoint0.10580.09550.05071.00001.00001.00003.3.3 MappingOnce the best possible configuration is determined for a unit cell in the ground structure, it is mapped to that region.

If there is a node from the unit-cell configuration that does not exist in the unit cell of the ground structure, it will be added using 3-D linear interpolation. After all the missing nodes are added, the unit-cell configuration can then be populated into the unit cell. The normalized stress values from step 4 of the augmented SMS method are scaled against the normalized diameter values from the unit cell library to determine the relative thickness of one strut to another.4 Design ExamplesIn order to validate the augmented SMS method, it will be applied to three design example. The first example is a 3-D curve cantilever beam. The second example is a micro aerial vehicle fuselage.

The first example is a 3-D curve cantilever beam. The second example is a micro aerial vehicle fuselage. These examples serve to validate the augmented SMS method and demonstrate its ability to design large-scale MSLS on complex-shaped parts with curve or non-rectangular surfaces.4.1 Curved Cantilever BeamThe first example is a simple, three-dimensional, curved cantilever beam with rectangular cross section. The beam is fixed at one end and has two point loads applied in the z-direction at the free end. The initial properties of this design problem are provided in Table 5. Ultimately, the primary goal of this example is to illustrate the modified algorithm in steps two, three, and four has allowed the augmented method to overcome the significant limitations with the previousimplementations of the SMS method which cannot be utilized in structures with curved surfaces.

For this example, the unit-cell regions do not have local coordinate systems that are the same as the global coordinate system. The weighting values from Eqn. 13 are set at: Wv = 1, Wvn = 1, Wp = 0. With these values, the topology was generated with 28 diagonal configurations. Other topologies were also generated by varying the weighting values. However, this topology has the best structural performance for this particular structure and loading condition. Figure 16 shows the final topology of the curved cantilever beam using active-set method, as well as the results from the intermediate steps. The strut diameter results are summarized in Table 6. The results show that the 28% method is able to converge about 6 times faster than either two-variable optimizations; however, the deflection result is worse than either two-variable approaches. The two 2-variable optimizations return identical results. However, the active-set method is able to converge faster.

The two 2-variable optimizations return identical results. However, the active-set method is able to converge faster. As mentioned in example 1, it is important to note that the diameter results reported for the SMS method are the actual diameters of the structure after being scaled against the solid body analysis and unit cell library scaling factors. The pre- scaling DMI(cid:28) and DMAX for the active set method are 0.4135 mm and 7.0704 mm, respectively.Table 5: Initial properties of the curve cantilever beam84.5 Outer Radius (mm) 74.5 Inner Radius (mm) 20 Width (mm) 105 Length(mm) 1 Loading Magnitude (N) Elastic Modulus (N/mm2) 1960 Unit-cell size in-plane (mm) 10 Unit-cell size out-plane (mm) 10 Target Volume (mm3)optimization approaches, a design space exploration/grid search was conducted. The design space exploration is done using pre-scaling values of diameters. This was performed by iterating both DMI(cid:28) and DMAX from 0.1 to 10 mm with increments of 0.1 mm.

The design space exploration is done using pre-scaling values of diameters. This was performed by iterating both DMI(cid:28) and DMAX from 0.1 to 10 mm with increments of 0.1 mm. The result is plotted in Figure 17. Based on the results from the initial exploration, a finer resolution of the design space was conducted around the apparent minimum by searching DMI(cid:28) from 0.3 to 0.5m and DMAX from 6.9 to 7.3 mm. with increments of 0.01 mm, as shown in Figure 18. The red diamond indicates the minimum found in design exploration. Diameter results that return the lowest objective function value are shown in Table 6. The pre- and post-scaling values from the design space exploration are included in two of the columns. These diameter values are found to be close to the values obtained from the SMS method.

The pre- and post-scaling values from the design space exploration are included in two of the columns. These diameter values are found to be close to the values obtained from the SMS method. Both two-variable optimizations were able to return results with much lesser design time than the design space exploration.Besidesthethree5000In addition to the augmented SMS method, ground truss approach was utilized to perform topology optimization for this selection scenario. Levenberg-Marquardt algorithm was used to optimize the ground truss. When compared with the ground truss approach, the augmented SMS has comparable performance result, but is able to converge much more quickly with about 70 times decrease in design time using two-variable approach and about 400 times decrease in design time using one-variable approach4.2 Micro Aerial Vehicle (MAV) FuselageThe final design example is a micro aerial vehicle (MAV) fuselage.

MAVs play a criticalrole in modern military operations as they allow easy surveillance in hazardous environment. The next generation of these aerial robotic systems needs to have enhanced take-off and landing capabilities, better endurance, and be adaptable to mission needs in varying conditions [28].

Interms of the design of the wings and fuselage of these MAVs, some types of structures and/or materials that are lighter, stronger and customizable are highly desired.WidthFigure 16: Final topology of curved cantilever beam using active-set methodOptimization ApproachTable 6: Optimization result for the curved cantilever beam SMS Design Space Exploration ScaledSMS Design Space Exploration (Pre-Scaled)SMS 28% AssumptionSMS Active- SetSMS Least-square MinimizationGround Truss LMDeflection(mm) Volume (mm3) Dmin (mm) Dmax (mm) Design Time (s)1.341 5501.0 1.13 2.43 7.80.901 5500.5 0.65 3.88 46.70.901 5500.5 0.65 3.88 52.80.900 5500 0.41 7.09 28600.900 5500 0.64 3.87 28600.780 5501 -- -- 3260In this design example, the fuselage is designed to withstand the impact when landing or crashing. There is a distributed load from the payload applied to the inner surface of the fuselage.

There is a distributed load from the payload applied to the inner surface of the fuselage. The weight of the motor and the tail are modeled as point loads at their centers of mass. This is done in ANSYS using a rigid link element. The equivalent couple is applied to the truss structure. All these weights are scaled by a factor of ten to simulate impact when crashing. The weight of the wing is small and assumed to have negligible contribution. A small area on the bottom of the fuselage is fixed to model the contact zone as the MAV is crashing. Multipleviews of the fuselage are shown in Figure 19 with key dimensions labeled.

A small area on the bottom of the fuselage is fixed to model the contact zone as the MAV is crashing. Multipleviews of the fuselage are shown in Figure 19 with key dimensions labeled. The initial properties of the design problem are provided in Table 7.Figure 17: Design space exploration for the curved cantilever beamFigure 18: Design space exploration with finer resolution around the solution for the curved cantilever beamThe objective is to achieve a target volume of 100,000 mm3 or 65% porosity relative solid model while minimizing the deflection of the structure. The ground structure for the fuselage is generated by TrussCreator with the desirable unit-cell sizes as provided in Table 7. There are 214 unit cells in the ground structure. The weighting values in Eqn. 13 set at: Wv = 0, Wnv = 1, Wp = 5. With these weighting values, the topology 100,00 crossed with is 0 configurations and 113 diagonal configurations.

The weighting values in Eqn. 13 set at: Wv = 0, Wnv = 1, Wp = 5. With these weighting values, the topology 100,00 crossed with is 0 configurations and 113 diagonal configurations. The topology matches the expectation based on the solid-body analysis. Other topologies were also generated by varying the weighting values. However, this topology has the best structural performance for this particular structure and loading condition.

Other topologies were also generated by varying the weighting values. However, this topology has the best structural performance for this particular structure and loading condition. The average displacement of selected nodes on the top of the fuselage, where the large displacement occurs according to the solid-body analysis, is used as the metric for deflection.Table 7: Initial properties for the fuselage 47 L (mm)A1 (mm)254theto90 Fmotor (N)A2 (mm)5.9A3 (mm)2.745 FTail (N)Fpayload (N/mm2)0.145D (mm)Unit-cell in-plane (mm) Unit-cell in-plane (mm) Unit-cell out-plane (mm)12 Modulus (N/mm2)196012 Poisson Ratio0.3Target Volume (mm3)8generated101The strut diameter results are summarized in Table 8. The active-set method returns the best deflection value but at the expense of design time with more than four the times fastest than assumption.

The active-set method returns the best deflection value but at the expense of design time with more than four the times fastest than assumption. method: However, 28% assumption returns the worst deflection result compared to the two 2-variable approaches. Of the 2-variable optimization methods, the least-squares minimization approachLA3DA2A1 (FRONT)Figure 19: Multiple views of the fuselagelonger28%is noticeable faster. The structure contains 101 crossed configurations and 113 diagonal configurations. The pre-scaled values of DMI(cid:28) and DMAX for the active set method are 0.5178 mm and 7.1406 mm, respectively.

The structure contains 101 crossed configurations and 113 diagonal configurations. The pre-scaled values of DMI(cid:28) and DMAX for the active set method are 0.5178 mm and 7.1406 mm, respectively. Figure 20 shows the final topology of the MAV design using the active-set method, as well as the results from the intermediate steps.FMotorFTailyFixed supportTailxFpayload(a) Side View(b) Bottom ViewFigure 20: Final topology for the fuselage.In addition to the three optimization approaches, design space exploration/grid search was conducted. The design space exploration is done using pre-scaling values of diameters. DMI(cid:28) and DMAX were iterated from 0.1 mm to 10 mm and with an increment of 0.1 mm. The increment, 0.1 mm, is coarse to reduce analysis time. However, even with the coarse increment, the design space exploration already takes a long time to complete. Exhaustive search is not a feasible solution for structures with large numbers of struts.

However, even with the coarse increment, the design space exploration already takes a long time to complete. Exhaustive search is not a feasible solution for structures with large numbers of struts. The result is plotted in Figure 21a. A finer resolution of the design space was conducted around the apparent minimum by searching DMI(cid:28) from 0.4 to 0.6 mm and DMAX from 7 to 7.3 mm with an increment of 0.01 mm. Figure 21b shows the design space zoomed into the region of interest. The red diamond indicates the minimum found in the design space exploration. Diameter results that return the lowest objective function value are shown in Table 8. The design space exploration diameter values are found to be close to the values obtained from the SMS method. Both two-variable optimizations were able to return results with much less design time than design space exploration.

Both two-variable optimizations were able to return results with much less design time than design space exploration. It can be seen fromFigure 21b, there is a valley in the design space exploration near the solution, which might cause the active set method to converge slowly because of the shallow gradient along one direction.Table 8: Optimization results for the fuselageOptimization Approach28% AssumptionActive- SetLeast-square MinimizationDesign Space Exploration (Pre- scaled)Design Space Exploration ScaledDeflection(mm)0.3270.2990.3190.2960.296Volume (mm3)100000.099973.0100010.0100000100000Dmin (mm)0.850.651.120.480.62Dmax (mm)3.044.163.297.284.22Design Time (s)378.61630.9508.976660766605. ConclusionsIn this article, two advances are reported for designing MSLS. First, computer-aided design technologies were developed for efficiently generating and representing MSLS.

ConclusionsIn this article, two advances are reported for designing MSLS. First, computer-aided design technologies were developed for efficiently generating and representing MSLS. More specifically, methods to construct conformal lattice structures were presented. Secondly, an augmented size matching and scaling design method for the design of conformal lattice structures was presented. This method enables us to design and efficiently optimize MSLS on complex-shaped parts by integrating the free-mesh approach in generating conformal lattice structures for the ground structure generation process. In addition, the method removes the need of rigorous large-scale multivariable topology optimization by utilizing a heuristic that reduces the multivariable optimization problem to a problem of only two variables, which combines solid-body analysis and predefined unit-cell library to generate the topology of the structure.

Based on this work, the following conclusions can be made:[a][b]Figure 21: [a] Design space exploration for fuselage, [b] Zoomed-in design space for the fuselageThe present CLS construction methods produce lattice structures on a wide variety of surface shapes, demonstrating the generality of the method. These methods produced CLS that was of high quality, since the unit cells were of nearly uniform size.In step 3 of the augmented SMS method, a new algorithm was developed to determine which unit cell of the ground structure the finite-element nodes fall into. After the finite element nodes are mapped to the correct unit cells in the ground structure, a stress transformation from global to local coordinate systems of the unit cells is conducted using standard rigid-body rotations.

Utilizing the property of second-order tensor, the stress was transformed to its proper local state to generate the correct topology as shown in the examples.For these particular structures with the given loading conditions, topologies generated with crossed configuration, diagonal configuration for Example 1, and a combination of cross and diagonal configurations for Example 2 return the best possible stiffness result.The weighting values Wv, Wvn, and Wp from Eqn. 13 should be set such that the topology is generated with the diagonal configuration for Example 1, and a combination of crossed and diagonal configurations in Example 2.

13 should be set such that the topology is generated with the diagonal configuration for Example 1, and a combination of crossed and diagonal configurations in Example 2. These weighting values are associated with Vσ∑ , Vnet, P∑ is a performance parameter that always favors the and cross configuration because it has the lowest strain energy.No generalized statement can be made on which unit-cell configuration should be selected because the method and the unit-cell library are only tested for a narrow set of design problems. • Between one-variable and two-variable optimizations, one-variable optimization using the 28% assumption consistently returns diameter results much faster than either two-variable optimization methods, but at the cost of structure stiffness. The two two-variable optimization methods produced very similar results in the first design example in terms of stiffness and design time.

The two two-variable optimization methods produced very similar results in the first design example in terms of stiffness and design time. However, least-squares minimization outperformed the active-set method in terms of design time in the MAV example. Due to the trade-off between design time and structural stiffness, the designer must make choose which design criteria are more important in his/her design when choosing the optimization approach.Overall, the augmented SMS method can be applied effectively in the design of conformal lattice structure with highly optimized stiffness and volume for complex surfaces. For simpler structures such as in Example 1, the augmented SMS method can reduce design time up to 70 times compared to normal topology optimization using a two-variable approach. In cases where topological optimization is infeasible, such as Example 2, the augmented SMS method can still effectively generate complex MSLS.

In cases where topological optimization is infeasible, such as Example 2, the augmented SMS method can still effectively generate complex MSLS. This approach removes the need for a rigorous topology optimization, which is a main bottleneck in designing MSLS. AcknowledgementThe authors gratefully acknowledge support from the US Air Force Research Laboratoryand Paramount Industries, Inc.References 1. Rosen, D.W., Computer-aided design for additive manufacturing of cellular structures. Computer-AidedDesign & Application, 2007. 4(5): p. 585-594.2. Gibson, I., D.W. Rosen, and B. Stucker, Additive Manufacturing Technologies: Rapid Prototyping to Direct Digital Manufacturing. 2010: Springer.3. Gibson, L.J. and M.F. Ashby, Cellular Solids: Structure and Properties. 1997, Cambridge: Cambridge University Press.4. Ashby, M.F., et al., Metals Foams: A Design Guide. 2000, Woburn, MA: Butterworth-Heinemann.5. Deshpande, V.S., N.A. Fleck, and M.F.

NFORMED-TO-CAD DESIGN SLOPED-EDGE ADAPTIVE SLICINGMohammad Hayasi, Bahram AsisbanpourIngram School of Engineering, 601 University Drive, Texas State University-San Marcos, 78666, Texas, USAAbstractMany mathematical and simulation based researches have shown the effectiveness ofadaptive and angular slicing and layering in minimizing errors in layered manufacturing processes and improving surface quality and build time. Recent technological advances such as 5-axis laser cutters, abrasive waterjets, and CNC machines have enabled physical part fabrication beyond typical vertical layered manufacturing with staircase errors. In this paper, a new curved- form adaptive slicing method is presented. In this method, multiple cuts of the edge boundary of each adaptive layer with variable cutting vector angles conform well to CAD model curved surfaces. The proposed system is compatible with 5-axis abrasive waterjet machines.

The proposed system is compatible with 5-axis abrasive waterjet machines. This system has been successfully tested on several virtual 3D CAD models.Keywords: Sloped-edge Adaptive Slicing, 5-Axis Cutting, Rapid Functional Part Fabrication, Rapid Tooling, Fully Dense Freeform Fabrication (FDFF)1. IntroductionRapid prototyping (RP) processes, such as direct metal laser sintering (DMLS), fused deposition modeling (FDM), selective laser sintering (SLS), stereolithography (SLA), and 3D printing have been in commercial use for decades across many business sectors. The rapid fabrication of a Computer-aided Design (CAD) model independent of any tooling is appealing as it would significantly reduce time, effort and cost - all crucial factors to the success of any business.

To build a rapid 3D prototype that geometrically matches its 3D CAD model, current commercial RP systems often apply very thin layers of metal foil, fine metal/ non-metal powder, or liquid resin laminating from the bottom to top of the CAD model such that steps appearing on the edges of the layered model become less significant. This, however, increases build time and thus fabrication cost. Thick-layer lamination with sloping surfaces offers a possible solution to overcome this problem by reducing the number of layers. Mechanical limitations of the commonly used RP systems make thick-layer building impossible. This is due to the fact that fusing, sintering, soldering, ultrasonic welding, and printing of the dense collection of particles and molecules in any selected bulk region of a variable thick layer is unachievable using present technology.

Methods of building functional metallic parts and tooling by means of abrasive waterjet and laser cutting of fully dense metal sheets in various thicknesses followed by alignment and attachment of layers have recently been developed (Nakagawa and Kunieda, 1984; Vouzelaud et al., 1992; Glozer and Bervick, 1992;Walczyk and Hardt ,1996; Bryden et al., 2001; Glen et al., 2011).The ability to rapidly cut various sheet metal thicknesses from small to large dimensionsusing abrasive waterjet or laser cutter machines helped to significantly reduce build time689compared to RP approaches that employed very thin layers of material. Fully dense rapid prototypes that are built up from non-uniform (adaptive) layers suffer from staircase errors appearing on the boundary edges of the layered CAD model (Figure 1).Layered boundarySignificant geometry distortion error (hatched area)CAD boundaryFigure 1.

Illustration of the magnitude of geometry deviation error using simple adaptive slicingThe sloped-edge adaptive slicing approach could improve the part quality to some extentbut it still results in a geometry distortion error (staircase) even on sloping edges (Figure 2).Layered boundaryReduced geometry distortion error (hatched area)CAD boundaryFigure 2. Illustration of the magnitude of geometry deviation error using sloped-edge adaptive slicingTo improve conformity to CAD model geometry, curved-form adaptive slicing on thesloped edges is presented. Multiple cuts of the edge boundary of each adaptive layer at variable cutting angles improve conformity to the CAD model curved surfaces (Figure 3).Layered boundarySignificant reduced geometry distortion error (hatched area)Thin layers that form thick adaptive layerCAD boundaryFigure 3.

Illustration of the magnitude of geometry deviation error using curved-form adaptive slicingAdaptive slicing for sloping surfaces takes advantage of applying thicker layers effectively to complex geometrical shapes containing such geometrical characteristics of general 3D models as convex and concave shapes, sudden changes of geometry, ruled surfaces, and sharp edges. In doing so, adaptive slicing, reduces the number of slabs and reduces geometric distortion. For successful implementation of such approaches, more axes of control are required, and the sloping surface path has to be obtained. Sloped-edge adaptive slicing proceeds through690two major steps: determination of the adaptive slice thickness by calculating various geometry deviation errors and connection of the corresponding upper and lower contour points of a layer in order to form the cutting trajectory. In an attempt to build parts with sloping surfaces in RP, Hope et al. (1997a) presented the TruSurf system.

In an attempt to build parts with sloping surfaces in RP, Hope et al. (1997a) presented the TruSurf system. This system first takes geometry and topology information from a sample CAD model file in Initial Graphics Exchange Specification (IGES) format. Accessing the B-spline surfaces of the CAD model and parametric data, the system seeks to determine the top and bottom contours resulting from the intersection of the XY plane with the B-spline surface at the defined Z level. Following the determination of the number of contour points defined by the user, the top and bottom contour points are computed by which the continuous connection of points by lines to approximate the actual CAD slices is facilitated. Using this algorithm and simple vector equations, the system succeeded in joining the corresponding top and bottom contour points to create cutting vectors and a trajectory.

Using this algorithm and simple vector equations, the system succeeded in joining the corresponding top and bottom contour points to create cutting vectors and a trajectory. The linear movement of the 5-axis machine head from start to end points of the connected contour lines and rotational movement of the machine head at two angles in cylindrical coordinates (cutting angle and polar angle) helped to build layers with slanting surface and variable thickness.Hope et al. (1997b) extended the application of the TruSurf system to the concise determination of variable layer thickness according to the surface geometry complexity. Also, they generated the layer sloping surface either inside the CAD boundary, outside, or a combination of both to meet the finishing requirements of the final prototype by comparing inside and outside tolerances with the computed cusp height.

For example, if finishing procedures are to remove material from the prototype surface, the tolerance would fall outside the CAD boundary (cusps of extra materials) resulting in outside inclined ruled surfaces (Figure 4(a)). Conversely if filling material needs to be applied to the surface of the prototype, then the tolerance would fall inside (cusps of missing material) and end up with inside sloped surfaces (Figure 4(b)). Either way requires material deposition or removal post-processing to make the prototype identical to its original CAD model.Missing materials (hatched areas)Beveled surfaceAdaptive layersCAD boundary(a)(b)Extra materials (hatched areas)Figure 4. Illustration of outside (a) and inside slicing position (b)Chamberlain et al. (1998) presented the zero integrated error ruled edge slice (ZIGERslice) algorithm in which cross sections of the sample 3D model are represented in STL format at one degree angle increments giving a total of 359 sections.

Sections are first obtained by averaging sample points of facet vertices. The connection of such sample points approximates the slice slope (cutting stream). Then, two important measures of error, average absolute error691and maximum absolute error, contribute to the determination of adaptive thickness slicing. In fact, the above errors were the negative and positive areas emerging between the CAD surface and approximate slope surface (Figure 5).Positive errorAdaptive layersCAD boundaryNegative errorApproximate beveled surfaceFigure 5. A view of ZIGER sliceUtilizing STL file format, Taylor et al. (2001) computed the intersection of a CAD modelrepresented in STL format with the Z plane at a specified Z value to create a chain of line segments forming the closed-loop or contour.

(2001) computed the intersection of a CAD modelrepresented in STL format with the Z plane at a specified Z value to create a chain of line segments forming the closed-loop or contour. The start and end points of these line segments are swept half the layer thickness above and below the cutting plane to form rectangles in the vertical or slanting depending on the angle of the cutting vector. These rectangles based on the start and end points of a line segment on the created contour and the cutting angle correspond to discreet cutter locations during machining of that contour (Figure 6).Swept half the above cutting planeSwept half the below cutting planeTrigger pointsFigure 6. An illustration of cutting trajectory formation.In conjunction with adaptive ruled layers, Koc and Lee (2002) took into account the generation of ruled surfaces obtained by joining the corresponding top and bottom contour points of an adaptive layer.

They concentrated on developing accurate connections of those points such that any twisted surfaces, surface approximation errors, and incorrect contour point joins would be avoided. The algorithm worked well for the connection of top and bottom contour points that were different in number and lent itself to the creation of ruled lines and surfaces, hence easing the generation of machine cutting paths suitable for 5-axis RP (Figure 7).692Ruled surfaceBottom contourTop contourFigure 7. Forming the ruled surface.In another attempt to form ruled surfaces between the contours of an adaptive slice,Gupta et al. (2004) proposed an algorithm to stitch together unequal optimized numbers of top and bottom contour points to obtain acceptable ruled lines, thereby generating well-shaped sloping surfaces and conforming closely to the part surface boundary.

Besides computing the optimum number of contour points and developing an adaptive slicing approach, they also provided a solution to the problem of cutting triangle patches that emerged from the multiple connection of one contour point with different points in its corresponding top or bottom contour. Yoo and Walczyk (2005) developed an advanced cutting trajectory algorithm for the profiled edge laminate (PEL) rapid tooling (RT) process (the process of the fabrication of large-scale dies and molds). Instead of slicing a CAD model represented by STL format, they directly performed the slicing process through SolidWorks API. A method for discretization of continuous profiles (i.e. the combination of arcs, splines, parabolas, and ellipsoids) of the CAD model has been developed.

A method for discretization of continuous profiles (i.e. the combination of arcs, splines, parabolas, and ellipsoids) of the CAD model has been developed. To stitch the obtained contour points appearing in front and back profiles for the purpose of accurate PEL surface reconstruction and the creation of optimized cutting vectors, an adaptive surface reconstruction algorithm consisting of Delaunay triangulation and a new complementary stitching algorithm were successfully implemented. The PEL technique is believed to be suitable for thick-layer rapid tooling that is suitable for 5-axis AWJ cutting machines leading to high quality prototypes.In a nutshell, despite the fact that many endeavors have been carried out in the scope ofbeveled-edge slicing, no research has been reported to address curved-form lamination for intricate beveled surfaces.

Therefore, the possibility of generating accurate machine path enabling curved-form sloped-edge lamination compatible with 5-axis abrasive waterjet (AWJ) cutters will enable fabrication of functional parts from fully dense sheets with geometrical characteristics very near to their original CAD models.Therefore, the main objective of the present research is to generate sloping surfaces onthe edge of an adaptive layer. The resulting surface has much less geometry deviation error and closely fits its CAD model boundary. The proposed system works in the following order: It takes in topology and geometry information from previously generated adaptive layers (Hayasi, Asiabanpour; 2011). Then, the thickness of an adaptive layer and the bottom and top contours of adjacent layers are fed into the proposed algorithm in the form of the continuous connection of vectors and their cutting vector with the facet normal for the creation of curved-form sloping surfaces.

Following curved-form adaptive slicing, a customized machine path and CNC code693compatible with 5-axis abrasive waterject cutters will be generated for any user-defined sheet thicknesses.The remainder of the paper is organized as follows: Section 2 briefly introduces the concept of rapid tooling (RT) and a new RT process named fully dense freeform fabrication (FDFF) with particular focus on metallic functional parts. In line with the quality improvement of the parts cut by the 5-axis abrasive waterjet, details of the new sloped-edge adaptive slicing system are presented in Section 3. Section 4 illustrates customized machine path and CNC-code generation which is consistent with current slicing systems and compatible with 5-axis abrasive waterjets without requiring machine setup modifications. Section 5 deals with implementation of the proposed system. Results achieved by application of the developed software on complex 3D CAD models are presented in Section 6.

Section 5 deals with implementation of the proposed system. Results achieved by application of the developed software on complex 3D CAD models are presented in Section 6. Finally, future work and conclusions are discussed in section 7.2. Rapid tooling and fully dense freeform fabrication processAlthough the proposed slicing algorithm is applicable to a variety of 5-axis processes thatallow variable thickness layering and slicing in different orientations, a brief review of the fully dense freeform fabrication (FDFF) process as the test bed and potential application of this algorithm would give the reader a better insight into the applicability of such an algorithm. The combination of variable thickness layering of any solid material, slicing in different orientations, and bulk layer attachment enables the FDFF process to build good quality complex parts from practically any dense material, including metals, plastic, wood, wax, ceramic, and even glass, at a very fast pace.

In this method, a CAD model is sliced into computer layers; the material sheet is then cut from the computer layers. Applying filler material between layers or direct joining of layers followed by layer compression mechanism and heating can be effective techniques to obtain acceptable layer attachment. The type of filler materials (e.g. glue, soldering or brazing filler depending on the applicability of filler material reacting to the heating source) and the acceptable distribution of compressive forces on all layers will be determining factors in attachment quality. Furnace, gas flame, or induction heaters are practical heat sources for heat based bonding methods (Figure 8).694Figure 8. Visualization of the sequences of the FDFF processDifferent aspects of the rapid tooling (RT) approaches have been addressed by many researchers such as Nakagawa and Kunieda (1984), Vouzelaud et al. (1992), Glozer and Bervick (1992), Walczyk and Hardt (1996), Bryden et al.

(1992), Glozer and Bervick (1992), Walczyk and Hardt (1996), Bryden et al. (2001), Gibbons et al., (2003), Himmer et al (2003), Perchtl et al. (2005), and Glen et al., 2011.3. Curved-form adaptive slicing3-1- Adaptive slicingThe input to curved-form adaptive slicing algorithms is an adaptive slice which carries a limited number of thin layers in order to be assigned an available thick layer. In adaptive slicing the CAD model is initially sliced at the maximum available thickness. Adaptive slicing then applies some specific tolerances to decide whether or not to break the thicker slab down into thinner ones. From this some basic information of a typical adaptive slice such as layer thickness and contour points of the top and the bottom slices of the associated layer are generated.

From this some basic information of a typical adaptive slice such as layer thickness and contour points of the top and the bottom slices of the associated layer are generated. The present system requires topology and geometry information of all thin layers that shape an adaptive slice with the desired thickness (Hayasi and Asiabanpour, 2011).3-2- Data structure formation for contoursFor each layer whether thin (consisting of a bottom and top slice) or thick (encompassingmore than two slices), an array of a specific data structure containing the required information for the slices (e.g. the internal and external closed-loop of each slice, the co-planar vectors of each contour, a surface identity assigned to each contour vector, and facet normal vector belonging to each contour vector) will be created.

The contours are obtained by the continued connection of vectors or line segments resulting from the intersection of the XY-plane with selected facets at the defined Z level in counter clock-wise (CCW) fashion. Since a contour vector or line segment is the result of connecting two points laid upon the edges of a facet, the695identity of a surface extracted from a facet will be readily obtainable by Autodesk Inventor API functions. Furthermore, a facet normal vector is added as an attribute into the data structure of each contour vector as a result of placing start and end points of a line segment on the edges of that facet. Each contour vector is a principle element in the determination of cutting angle and trajectory (Figure 9).Internal contourContour vectorExternal closed loop/contourCCWData Structure (Contour): 1. Start Point (SP) 2. End Point (EP) 3. Contour Vector (𝑉⃗ ) 4. Facet normal (𝑁⃗⃗ ) 5. Slice No. (i) 6. Contour No.

Consequently, a cutting vector ( ⃗⃗⃗⃗ ) can be computed by taking the cross product of the obtained contour vector and a facet normal ( ⃗⃗⃗⃗⃗ ) that had been added into the vector record in the defined data structure. Since all contour vectors are ordered in counter clock-wise manner, for internal closed-loops (contours) the cross product of696vectors must be given by ⃗⃗⃗⃗ ⃗⃗⃗⃗⃗ ⃗⃗⃗⃗⃗ and by the opposite ( ⃗⃗⃗⃗ ⃗⃗⃗⃗⃗ ⃗⃗⃗⃗⃗ ) for external contours in order to lead the pick of the cutting vector in the direction of the cutting stream. Following calculation of the above vectors, the two rotational and cutting angles are obtained as follows:Cutting Vector angle:⃗⃗⃗⃗⃗ ‖ ‖‖ ⃗⃗⃗⃗⃗ ‖Rotational angle: ⃗ ⃗⃗⃗⃗⃗ ‖ ⃗ ‖‖ ⃗⃗⃗⃗⃗ ‖(1)Figure 11 is a graphic view of the vector computation on the simple adaptive sliceincluding a few thin slices with multiple loops within a thin slice.(a)(b)Figure 11.

A 3D view of CAD model for vector computation (a) visual representation of machine head angles (A,C) and cutting vector (b)3-3- Cut path difference for convex and concave curvesFor cutting curve edges, the proposed system should be capable of recognizing concaveor convex curves throughout the surface of the layered model while tracing all contour vectors in a CCW direction. This is an important part of the system as it will help to determine the appropriate action to cut the curved regions. For convex edges, the system will provide a custom machine path along which the multiple cuts required for that edge will be possible. Using multiple cutting vectors for a single convex edge results in a curved-form edge closer to its original CAD edge with minimal geometric deviation (Figure. 12 (a)). For a concave edge, it is impossible to apply multiple cuts in this fashion, as doing so results in removal of excessive material in the first single cut (Figure.

12 (a)). For a concave edge, it is impossible to apply multiple cuts in this fashion, as doing so results in removal of excessive material in the first single cut (Figure. 12(b))6976 cutting vectors based on the number of thin slices32112(a)3(b)Proper cutting vector45 5 6Improper cutting vector (leads to missing material)456Figure 12. Two cutting strategies: (a) Multiple cuts to form a convex edge (b) single cut for a concave edgeAs a result, the top and bottom points of the concave curve and subsequently theconnection of these points need to be identified in order to create a single cutting vector, as a guide for the AWJ cutting stream.3-4- Convex and concave curve identificationIn conjunction with the identification of convex and concave edges, a contour vector at aparticular slice number, for example, k, is first picked and its start and end points ( ) and its cutting vector ( ⃗⃗⃗⃗⃗ ) is extracted from the data structure.

Then, the midpoint ( ) of is found and mapped onto a plane parallel to the XY-plane and passing through slice number 1. To find a second point that lies on either the concave or convex curve and located on ⃗⃗⃗⃗⃗⃗⃗⃗ ) tangent to the surface at point one of the contour vectors of the bottommost slice, a vector ( is projected on the same plane as the midpoint and a line parallel to that projected vector ⃗⃗⃗⃗⃗⃗⃗⃗ ) and passing through the projected midpoint( ) is obtained. The points resulting from ( the intersection of the line (L) with a short list of line segments (lines that surely intersect with the line L) located on the first slice are obtained. Among the points there is a point ( ) placed on a contour vector whose surface identity equals that of ⃗⃗⃗⃗⃗ and also lies on a vector (forming by the candidate point and point ) co-directional with obtained, a vector passing through that point and tangent to the part surface is created and its angle ( ) with ⃗⃗⃗ is computed.

Again at point , a vector tangent to the part surface is created and subsequently an angle ( ) between this vector and ⃗⃗⃗ is found. A curve edge is regarded as convex if is greater than otherwise the edge is concave (Figure 13).⃗⃗⃗⃗⃗⃗⃗⃗ . Once the desired point is698Figure 13. Graphical representation of how a convex edge is identified3-5- Curved-form adaptive slicing main algorithmOnce the adaptive slicing from the bottom to top of the CAD model is accurately accomplished, an external contour for each adaptive layer containing a number of thin slices is selected from among contours in the top slice of the layer. The initial vector in the chain of the vectors is obtained from the data structure of the contour vectors belonging to the selected contour and the possibility of being convex or concave at this point is evaluated using the method explained above.

With regard to concave curves a straight sloping surface is formed by mapping the start and end points of the selected vector onto a plane parallel to the XY-plane and accommodating the contours of the bottom slice. Then two lines passing through the mapped points and parallel with the projected cutting vector of the top contour vector are drawn. The intersection of the lines with the corresponding line segments (contour vectors) on the bottom slice results in two additional points. Eventually, the desired beveled surface, which is created by the connection of those four points, would be a useful tool for both visual simulations of the curved-form cutting and as a source to generate 5-axis AWJ cutter machine trajectories (Figure 14(a)).Unlike for concave curves the system is required to provide multiple slanting surfaces such that its output conforms to the curvature of convex region.

With the assumption that the curvature of the part surface at the initial selected contour vector is found to be convex using the method explained, two lines passing through the start and end point of that vector and parallel to the cutting vector at this location are created. The two lines then intersect with a plane parallel to the XY-plane at a Z-level where the bottom slice is located (Figure 14(b)). The points resulting from the line-plane intersection are simply obtained by the following equation 2.Line equation:, ,(2)By settingthen699(a)(b)Figure 14. Illustration of a single cutting stream for concave edge (a) and multiple cutting streams for convex edge (b) formed by a number of thin slicesThe first slanting surface formed by the connection of the four points is retained in computer memory for further use in machine path generation.

To complete beveled surface formation on the top slice, the system seeks to trace all the vectors within either external or internal loops of the slice and apply the above described algorithm. For the remainder of the thin slices, if the curvature of a part surface at a certain vector is found to be concave, the surface formation must be neglected based on the rule defined for such concave curves (i.e. single sloped cutting surface on concave edges). An overview of the developed algorithm presented as a flowchart is shown in Figure 15.700Figure 15. Curved-form adaptive slicing algorithm4. Machine path and CNC-code generationCurrently available CAD/CAM and CAM systems are not capable of generating code forcurved-form 5-axis AWJ multiple cutting paths. Thus some custom modification of G-codes compatible with 5-axis AWJ machines would enable the production of objects very nearly resembling their CAD designs.

Thus some custom modification of G-codes compatible with 5-axis AWJ machines would enable the production of objects very nearly resembling their CAD designs. Since a CAD model is mainly layered in adaptive manner by the use of available sheets on the shop-floor, the size of such sheets must be an input to fit different adaptive slices in their associated sheets according to the sheet thickness. Following CAD model lamination, a group of slices whose thickness falls in the predefined sheet thickness are put together in matrix fashion on the sheet. In order to cut the layers in sequence, the position of each layer on its corresponding sheet is appropriately numbered. The CNC codes are written in such a way that an AWJ machine head would start cutting the layers from the thin through the thick sheet (Figure 16).701Figure 16.

The CNC codes are written in such a way that an AWJ machine head would start cutting the layers from the thin through the thick sheet (Figure 16).701Figure 16. Top view of the arrangement of 7 distinct adaptive slices onto two available sheets with specified thicknessesIn G-code generation, there are four basic parameters required for the purpose of 5-axiscutting, namely, linear movement of the machine head from start to end point of each contour vector, rotational and cutting angle of the machine head, and machine feed rate. The first four input data (SP, EP, , ) were simply obtained while generating the adaptive slices and applying the foregoing algorithm to each thick layer. There have been a few research efforts to determine how to use machine parameters and cutting depth to assign feed rate to get a good surface roughness quality.

There have been a few research efforts to determine how to use machine parameters and cutting depth to assign feed rate to get a good surface roughness quality. As a result, this part of the problem is not in the domain of our research and we mainly refer to previous work that arrived at a comprehensive feed rate formula. The following formula is adopted from the research work of Limbachiya and Patel (2011).()√(3)Where,= orifice diameter, = insert diameter, R= ratio of mass flow rate of abrasive to flow rate of water, = density of water, = specific energy of material, p= pressure in bar, and =cut depth.Parameters ( , , R, , , and p) are easily obtained, but to calculate we must measure the depth of either vertical or slanting surfaces at every contour vector position on the adaptive layer. The calculation of is graphically depicted in Figure 17.702Figure 17. The explicit computation of in the projected view of a sample adaptive slice5.

The calculation of is graphically depicted in Figure 17.702Figure 17. The explicit computation of in the projected view of a sample adaptive slice5. System implementationThe algorithm was implemented using Visual Basic for Application (VBA) and AutodeskInventor mechanical design software. Autodesk Inventor API functions were used to gain direct access to topology and geometry information of any complex CAD model. A user interface form facilitates data entries (i.e., slice direction in either X, Y, or Z directions, various sheet thicknesses, allowable errors for boundary area deviation, triangle area, number of available sheets, sheets thickness, and the width and length of the sheets). Figure 18 illustrates the user interface form for triple slicing approaches that was developed in response to a user request. Even though the generated 5-axis AWJ machine codes are saved in one text file, the software output can be obtained layer by layer in the 2D and 3D versions.

Even though the generated 5-axis AWJ machine codes are saved in one text file, the software output can be obtained layer by layer in the 2D and 3D versions. In the 2D version, the generated layers containing line segments and sloping surfaces are grouped according to their sheet thickness and spread on the sheet rectangle surfaces in matrix form followed by numbering each adaptive slice. In contrast, in the 3D version all layers along with their cutting surfaces in their own thickness are stacked together in the build direction to form an approximate real 3D solid model for better visualization of the final physical part being fabricated by the FDFF process.703Figure 18. The user interface form.6. System results, verification, and validationThe implemented system was successfully tested for slicing a variety of models with diverse types of surfaces (i.e., flat, convex, and concave), different slicing direction, and different number of sheets with different thicknesses.

The following example along with different stages of the curved form adaptive slicing is presented below:The entries for model were: sheet thicknesses and sizes T1: 0.1 in, L1: 20 in, W1: 15 in,T2: 0.25 in, L2: 15 in, W2: 30 in, and Cut direction: Z (Figure 19)704Figure 19. (a) 3D CAD model and (b-d) 3D and side view of curved-form adaptive slicing7- ConclusionIn this article a new curved-form adaptive slicing method is presented. In this method, multiple cuts of the edge boundary of each adaptive layer with variable cutting vectors angles conform well to the CAD model curved surfaces. The results showed that the staircase effect, a common error in adaptive slicing, existing on either concave or convex edges has been significantly reduced utilizing this method. The proposed system is compatible with 5-axis abrasive waterjet machines. This system has been successfully tested on several models. The main application of the proposed system is in the FDFF process.

This system has been successfully tested on several models. The main application of the proposed system is in the FDFF process. Metallic sheets with variety of predefined thicknesses can be cut with angular direction by a 5-axis AWJ machine and according to the tool path generated by the proposed system. Then, layers are aligned and bonded in the FDFF process. The bonding process is typically by the induction heater and by using the brazing materials. Therefore, the mechanical properties of the fabricated parts are good enough to be used as the functional parts. The FDFF process has been extensively used and tested for the parts cut by a 2 axis AWJ machine. The future of this research is the test of the FDFF process for the parts cut by a 5 axis AWJ machine, subject to purchase/access.ReferencesAsiabanpour, B., & Khoshnevis, B. (2004).

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperConsiderations of internal void generation process by observation of melting and solidification behavior in low temperature laser sintering of PEEKT. KIGURE*, Y. YAMAUCHI*, T. NIINO†Tokyo Metropolitan Industrial Technology Research Institute †Institute of Industrial Science, the University of TokyoAbstractLaser sintering of PEEK performed below the crystallization temperature has been achieved in previous study by low temperature process that anchors the part to a rigid base plate to suppress warpage during processing. However, significant surface roughness and large internal voids are sometimes generated in the parts built by low temperature process, and there are problems in stability of parts quality.

However, significant surface roughness and large internal voids are sometimes generated in the parts built by low temperature process, and there are problems in stability of parts quality. The purpose of this study is to contribute to quality improvement of laser sintering of PEEK by low temperature process. It was attempted that clarify the process of surface roughness and void generation by observing the melting and solidification behavior of the material during process with a video camera. From these observation results, it was assumed that the amount of volume change from powder to liquid due to melting and the amount of shrinkage due to solidification affect part quality such as surface roughness and internal voids.IntroductionPolyEtherEtherKetone (PEEK) which is one of the high-performance plastics with high strength, high chemical, and heat resistance, can be used as a feedstock material for laser sintering.

General laser sintering process of PEEK requires a powder bed temperature about 300 °C to suppress parts warpage during the process [1-3]. Since 300°C is above the melting temperature of most polymer materials, laser sintering of PEEK is limited to high heat resistant machine. Therefore, enabling laser sintering of PEEK by ordinary machine that does not have high heat resistant is one of challenges. The authors have been proposing low temperature process as a way to achieve laser sintering of PEEK on ordinary machines that are not highly heat resistant.Low temperature process is the method of anchoring the part to a rigid base plate instead of maintaining the powder bed temperature between the melting point and the recrystallization temperature to suppress warpage during the process. This method can set powder bed temperature lower than recrystallization temperature.

This method can set powder bed temperature lower than recrystallization temperature. In previous study, it has been shown that setting the powder bed temperature lower than the recrystallization temperature improves the material recycling rate [4], and parts with various crystallinity can be obtained in various powder bed temperatures [5].The authors have succeeded that laser sintering of PEEK at a powder bed temperature of 170 °C that is the same as the laser sintering process for Polyamide12, the most common material for laser sintering by using a low temperature process [6]. However, obtained PEEK parts have had some problems, such as significant rough surface and large diagonal voids [7]. Since voids and131rough surfaces cause a reduction in part strength, it is necessary to clarify the mechanisms of voids and rough surface generation.

Since voids and131rough surfaces cause a reduction in part strength, it is necessary to clarify the mechanisms of voids and rough surface generation. Since it is obvious that generation of rough surface and inner voids are related to the flow behavior of molten polymer powder, in-situ observation is effective. Many in situ observations have been reported for powder bed fusion of metals. The shape of the melt pool, the balling phenomenon, and the generation of fumes due to excessive energy supply have been observed, and the mechanism has been considered [8-10]. On the other hand, report of in-situ observation in powder bed fusion of polymer is few. Gardner et al. used optical coherence tomography to investigate melt pool shape [11] and Klamart et al. showed curling during process by using thermal imaging [12]. Melvin et al. used high-speed microscope and showed vapor bubble formation and entrapment in liquid pool [13]. Sassaman et al.

showed curling during process by using thermal imaging [12]. Melvin et al. used high-speed microscope and showed vapor bubble formation and entrapment in liquid pool [13]. Sassaman et al. designed laser sintering machine implemented heat resistant high-speed microscope [14]. In this report, it has been found that laser melting of the polymer powder occurs with a delay from the passage of the laser spot and over extended time.In this study, it is attempted that clarify the process of surface roughness and void generation by observing the melting and solidification behavior of the material during process with a video camera. In addition to observing the parts building process by low-temperature process, the tendency of surface roughness to occur is investigated by laser irradiating different powder layer thicknesses.

In addition to observing the parts building process by low-temperature process, the tendency of surface roughness to occur is investigated by laser irradiating different powder layer thicknesses. This study is to contribute to quality improvement of laser sintering of PEEK by low temperature process.MethodologyIn this study, PEEK is used as material and its melting and solidification behavior is observed with video camera. Not only the behavior during low-temperature processes, but also the behavior at various powder thicknesses is observed for understanding the mechanism of rough surface and void generation.Material and Machine The material used in this study is PEEK powder (Vestakeep 2000FP, Evonik inc.). Since this powder is not intended for use in additive manufacturing and contains large particles, the powder was sieved with mesh size of 106 μm to remove large particles. Average particle size after sieving was 66 μm. Table 1 shows the material properties.

Average particle size after sieving was 66 μm. Table 1 shows the material properties. Thermal properties were determined by differential scanning calorimetry, and densities were obtained from density measurement tests. Material with no thermal history was used.Glass transition temperature Approx. 150℃Table 1. Material properties. Melting pointRecrystallization temperature 299 ℃343 ℃Bulk density0.3 g/cm3Ture density1.3 g/cm3The machine that modified commercial laser sintering machine (RaFaEl 300C®, Aspect inc.) was used. This machine is equipped with a CO2 laser and laser spot dimeter can be changed from 350 μm to 1600 μm using defocusing function.132Test Condition of low temperature process In this study, specimen build tests and single layer tests for various powder thicknesses were performed. Table 2 shows the scan parameters. The powder bed temperature of 170°C is lower than the recrystallization temperature of PEEK.

Table 2 shows the scan parameters. The powder bed temperature of 170°C is lower than the recrystallization temperature of PEEK. PEEK plate prepared by injection molding were used as base plates for both moldings. The same laser parameters sets and baseplate were used for both experiments.Spot diameters 1685 μmLaser power 40 WTable 2. Scan parameters sets. Powder bed temperature 170 ℃Scan spacing 0.02 mmScan speed 5 m/sIIScan strategy LasterScan direction SingleIn the specimen build test, the specimen shape was half size tensile specimen. Layer thickness was set at 0.1mm and laser scanning direction was the longitudinal direction of specimen as shown Figure 1. Specimen was built on the support structure.Laser scan path------------------------ -Support structureFigure 1. Specimen shape and scan direction of specimen build test.Figure 2 shows a schematic image of the single layer test for various powder thickness.

Specimen shape and scan direction of specimen build test.Figure 2 shows a schematic image of the single layer test for various powder thickness. The powder thicknesses were 0.1 mm, 0.2 mm, 0.3 mm, 0.4 mm, and 0.5 mm. The thickness of the powder was controlled by the layering pitch. Scanning shape was 80mm × 10mm rectangle and scanning direction was the longitudinal direction of the shape.Itrndiate laser at fast powder Spread powder for next thickness (O. lnun)powder thickness (layer thickness 0.1mm)ltrndiate laser at next powder thicknessFinish laser iirndiation to vruious powder thicknessFigure 2. Schematic image of the single layer test for various powder thickness.Relationship between powder thickness and layer thicknessAs reported by Childs et al., the powder thickness that needs to be melted in the lasersintering process is thicker than the layer thickness[15].

The state of the material changes from powder to block through melting and solidification, which also changes the bulk density. The relationship between powder thickness (𝑡𝑝𝑜𝑤𝑑𝑒𝑟) and layer thickness (𝑡𝑙𝑎𝑦𝑒𝑟 ) is showed as equation (1). Here, block density (𝜌𝑏𝑙𝑜𝑐𝑘) (≅ true density), powder density (𝜌𝑝𝑜𝑤𝑑𝑒𝑟) are, respectively.133𝑡𝑝𝑜𝑤𝑑𝑒𝑟 = 𝑡𝑙𝑎𝑦𝑒𝑟𝜌𝑏𝑙𝑜𝑐𝑘 𝜌𝑝𝑜𝑤𝑑𝑒𝑟Observation method during process Figure 3 shows observation set up. Digital video camera (CANON inc.) with 2.07 million pixels and frame rate of 30 fps was used. A camera was placed outside the machine and the melting and solidification behavior was monitored through the machine window.Figure 3. Overview of the observation experiment.Experimental resultFigure 4 shows a tensile specimen (top) and its cross section (bottom). It is showed that the surface of the specimen was significantly roughened like a wave. The wave spacing was more than 1 mm, which was larger than the scan spacing.

It is showed that the surface of the specimen was significantly roughened like a wave. The wave spacing was more than 1 mm, which was larger than the scan spacing. The cracks extended diagonally from the surface across each layer, and the slopes of the cracks were tilted in the laser sweep direction.p e e w sr e s a Ln o i t c e r i dLaser scanning direction10 mmLaser sweep direction1 mmFigure 4. Specimen built by low temperature process. Top is the whole specimen, and bottom is cross section of specimen.134(1)Figure 5 shows the state of laser irradiation during specimen build test. This picture shows how the powder near the boundary between the melted and unmelted powder was drawn into the melted region like an "avalanche" during laser scanning. This avalanche phenomena occurred periodically during laser scanning.Sweep directionFigure 5.

This avalanche phenomena occurred periodically during laser scanning.Sweep directionFigure 5. Appearance of avalanche.Figure 6 shows the state of the solidification process from the end of laser irradiation in 10 seconds interval and its schematic image. At the end of laser irradiation (0 second), solidification has already started on the left side of the specimen, which is the beginning of irradiation. 10 seconds after the end of irradiation, the right side of the specimen, which is the last irradiated area, began to solidify, but the center of the specimen had not yet solidified. Then, 20 seconds after the end of irradiation, solidification was completed.t = 0 [s]t = 10[s]t = 20[s]0 [s]10 [s]20 [s]Dsolidified areaFigure 6. Solidification process of specimen. Top is photo during laser irradiation, bottom is schematic image.135Figure 7 shows the results of single-layer tests for various powder thicknesses.

Solidification process of specimen. Top is photo during laser irradiation, bottom is schematic image.135Figure 7 shows the results of single-layer tests for various powder thicknesses. Waveform rough surface and thermal decomposition were observed at powder thicknesses of 0.4 mm and 0.5 mm. Figure 8 shows the state of melting behavior at each powder thickness. Avalanche phenomena were observed at powder thicknesses of 0.4mm and 0.5mm where waveform roughness occurred.0.1mm0.2mm0.3mm0.4mm0.5mmFigure 7. Result of difference powder thickness melting test. Thermal decomposition and rough surface was observed in thickness 0.4mm and 0.5mm.0.1mm0.2mm0.3mm0.4mm0.5mmFigure 8. Behavior of each powder thickness during laser irradiation.

Thermal decomposition and rough surface was observed in thickness 0.4mm and 0.5mm.0.1mm0.2mm0.3mm0.4mm0.5mmFigure 8. Behavior of each powder thickness during laser irradiation. Thermal decomposition smoke generated in thickness 0.4mm and 0.5mm and avalanche was observed in thickness 0.5mm.DiscussionThe cross section shown in Figure 4 shows that diagonal voids were generated from the rough waveform surface. The period of the waveform was larger than the scan spacing and did not occur with each laser scan. The behavior shown in Figure 5 confirms the separation of the molten region due to avalanche that occurred during laser scanning.

The behavior shown in Figure 5 confirms the separation of the molten region due to avalanche that occurred during laser scanning. Therefore, the waveform surface was formed by periodic powder avalanches separating the molten region, and it is assumed that the occurrence of diagonal voids was the effect of the remaining internal waveform surface irregularities.The solidification behavior was not completely dependent on order of laser irradiation, region of shorter laser scanning length was solidified faster than the region of longer scanning length. This is related to the cooling rate of each part, and it is indicated that the solidification behavior of low temperature process has a shape dependence. In addition, even at a powder bed temperature that was more than 100 °C lower than the recrystallization temperature, it took about 20 seconds from the end of laser irradiation to complete solidification.

This was due to the low thermal conductivity of the powder and heating from the already formed underlying layer.136Solidified surface was rougher as powder layer thickness increasing. Especially, waveform surface appeared at a powder thickness of more than 0.3 mm. This is because the movable range of the melted polymer increased as the thickness of the powder layer increased, causing polymer flocculation and flow. The reason for the thermal decomposition in the thick powder layer is assumed as follows. In thin powder layer, the surface temperature does not increase because the polymer adhered to the base plate as soon as it melted and dissipated heat to the base plate.

In thin powder layer, the surface temperature does not increase because the polymer adhered to the base plate as soon as it melted and dissipated heat to the base plate. On the other hand, in the case of a thick powder layer, the melted polymer flowed downward and then adhered to the base plate, so it is assumed that the surface temperature increased as more heat was received from the laser irradiation before it was diffused to the base plate.In considering to specimen building test, due to the powder thickness that is melted is about 0.4 mm in the case of layer thickness of 0.1 mm from equation (1), the avalanche has occurred and generated significant rough surface. In addition, diagonal void is generated from this rough surface as starting point. From these results, it is assumed that powder thickness is needed to keep below 0.3mm to avoid from rough surface and large inner voids.

From these results, it is assumed that powder thickness is needed to keep below 0.3mm to avoid from rough surface and large inner voids. To reduce the thickness of the powder layer, it is effective to reduce the layer thickness or increase the bulk density of the powder. In the case of this study, from equation (1), It is assumed that lamination layer thickness must be below 0.07 mm or bulk density of PEEK powder must be over 0.43 g/cm3 to achieve below 0.3 mm powder thickness that is melted.It is assumed that the avalanche phenomenon, which causes surface roughness and internal voids, is affected not only by polymer fluidity, but also by temperature distribution and shrinkage during transition from powder to block.

In the future, more detailed investigations such as high speed camera observations and thermal analysis are required to understand and resolve this phenomenon.ConclusionThis study attempted observation of melting and solidification behavior in low temperature process by using digital video camera. As a result, it was found that avalanche, which occurs during laser scanning with thick powder layers, is the cause of the large rough surface. To suppress this phenomenon, the thickness of the powder layer, which is different from the layering pitch, must be kept at a value that does not cause avalanche. The knowledge from this study is not only contribute to the quality improvement of low temperature process, but also provide important information regarding the requirements for the materials.AcknowledgmentsThis research is a part of “R&D Support Project for Growing SMEs” supported Ministry of Economy, Trade and Industry, Government of Japan.Reference[1] N.

For this, a machine- and process-independent rating of the processability is necessary. The established standards for the tensile strength test apply to specimens with nearly isotropic mechanical properties. The FDM process generates anisotropic parts. The properties are mainly influenced by the machine quality and the data processing. It is not possible to test a material for FDM independently of the machine and the data processing. In this paper, machine and process- specific influences are investigated. Considering these influences, a custom-built specimen is created to test the tensile strength of the welding seams for polyamide 6. This procedure allows a machine- and process-independent rating of the processability in terms of tensile strength for different materials.IntroductionThe Fused Deposition Modeling (FDM) process is one of the most commonly used additive manufacturing processes.

In this method, the semi-finished product, a filament of a thermoplastic polymer, is melted and forced through a nozzle. The continuous positioning of this nozzle allows the polymer to weld together strand by strand and layer by layer to produce a component. The energy for the welding of the individual strands largely results from the thermal energy of the deposited polymer melt and, if present, through a heated chamber.It is desirable to be able to use a similarly wide variety of materials with this method as, for example, in the profile extrusion or injection molding technology. Therefore, the processing suitability of any thermoplastic polymer should be estimated based on the material properties or characteristics in advance of the processing. This is currently not possible because, in contrast to conventional methods, there is a lack of knowledge about the required and desirable material properties for processing by the FDM process.

This is currently not possible because, in contrast to conventional methods, there is a lack of knowledge about the required and desirable material properties for processing by the FDM process. By gaining a better understanding of the process, a knowledge base should be created to increase the variety of materials that are available. This work is conducted in cooperation with the company ALBIS PLASTIC GmbH and under the NRW “Fortschrittskolleg Lightweight – Efficient – Mobile” (FK LEM).With the aim of material development, this paper sets out to identify test methods and test specimens with the aid of which the suitability of various materials for the FDM process can be quantified. To evaluate the suitability, characteristic data are defined. The aim here is to be able to470test the effects of modified material properties, isolated from other influences. The machine quality and data processing should, wherever possible, not have any effect on the results.

The machine quality and data processing should, wherever possible, not have any effect on the results. In order to enable the selection and/or development of such specimens and test methods, the FDM process must be adequately understood. Any influences that occur must be identified and taken into account when designing the test methods and test specimens.In this paper, process-specific influencing factors on the product are described. Subsequently, a specimen is designed and the corresponding test method is presented for assessing the weld seam strength. Based on this, the influence on the weld seam strength of a change in viscosity in fused deposition modeling is tested and evaluated for PA 6. The materials and the determined material properties have been supplied by our research partner ABIS Plastic.Influencing factors in FDMIn the FDM process – as in every manufacturing process – there are certain influencing factors that affect the production result.

Through these and the process itself, production-specific structures are formed in the part. The influencing factors can be divided up according to [Klo16, p. 10–14] into material-related, process-related and machine-related factors. The following chapter discusses the main influencing factors, orientated to this subdivision.Influences from the material and the semi-finished productProcess-specific demands are made on the material used in the FDM process and also on the plastic filament used as the semi-finished product. The influences exerted by the material or material properties on the production result in the FDM process are at present not sufficiently well- known and need to be investigated. A sufficient understanding of the process and experience from conventional plastics processing processes serve to derive fundamental preconditions for the materials used and the properties of the semi-finished product.

For example, in the classic FDM process, the material is converted into the molten state and fused. This means that only a thermoplastic material can be used.No dirt particles must be present in the semi-finished product because of the specified part quality and the diameter of the nozzle used in the FDM process. Large dirt particles can block the nozzle and bring the process to a halt. It must also be ensured that the material is dry enough during processing. Any moisture can outgas during processing, leading to pressure peaks in the plasticizing unit, uncontrolled delivery of the plastic melt and foaming of the material. The semi- finished product must also have a constant cross-sectional area over its entire length. The reason for this is that the amount of plastic delivered in the FDM process is calculated via the controlled feed distance of the semi-finished product and the cross-sectional area contained in the software.

If the actual cross-sectional area deviates from that in the software, the amount of plastic melt delivered in the process will not agree with the calculated quantity of delivered plastic and the part will be faulty. For the same reason, but also to avoid bubbles and blisters in the discharged plastic melt and defects in the final part, the semi-finished product must not contain any pores or voids.The semi-finished product used in this research project is produced under laboratory conditions from fine pellets, taking account of the previously derived requirements. The quality of the semi-finished product is monitored via a visual examination of the transparent melt during471processing and of the extruded product. Furthermore, the diameter and ovality are controlled by an inline laser measuring system of type Accuscan 5010 from the firm BETA LaserMike with several laser curtains.

Furthermore, the diameter and ovality are controlled by an inline laser measuring system of type Accuscan 5010 from the firm BETA LaserMike with several laser curtains. Adequate drying is carried out before extrusion into the FDM semi-finished product and FDM processing.Influences from the process/data processingThe building of a part through the laying of strands next to and on top of one another as well as the sequence, direction and method of strand deposition have an influence on the production result. The manufacturing process is defined by the process principle and the set production parameters. The translational accelerations and velocities of the tool that have to be reached in the process as well as the melting performance of the plasticizing unit are subject to machine-specific limitations. Influences that are process or machine-related are sometimes directly connected.

Influences that are process or machine-related are sometimes directly connected. With an adequate understanding of the FDM process, factors influencing the production result are easier to comprehend. The following section will therefore briefly explain the principle of action of the FDM process and the various process phases. After this, a list will be given of important influencing factors.Principle of action and process phases of Fused Deposition ModelingThe principle of action of FDM can be basically described by two functions. One function is the translational motion of the tool. The other function is the delivery of a flow of plastic melt by the tool (nozzle).

One function is the translational motion of the tool. The other function is the delivery of a flow of plastic melt by the tool (nozzle). The strand-shaped deposition of the plastic that is typical of the FDM process is attained by overlapping, in the plane, a translational motion of the tool while simultaneously delivering plastic melt.One way of differentiating between various process phases in fused deposition modeling comes from the type of strand deposition. This can be interrupted during the process or carried out non-stationarily or stationarily (cf. [Bel02, p. 34-35, 147]). During an interruption to strand deposition, no plastic melt is deposited as a strand. Translational motions can be performed without the delivery of plastic. These are necessary in the FDM process to jump, for example, to another section of the part, or to begin a new layer. During strand deposition, a distinction can also be made between two other process phases.

During strand deposition, a distinction can also be made between two other process phases. Non-stationary strand deposition is characterized by acceleration processes of the tool and/or the plastic delivery. Acceleration processes are necessary because of the existing mass inertia, e.g. with changes of direction. In addition to the translational motion velocity of the tool, the velocity of the plastic delivery can also be adjusted. At present, various techniques are being used and investigated in order to obtain a strand deposition that is as constant as possible despite changing translational tool velocities (cf. [GRS16, p. 154], [Bel02, p. 34], [HJ07, p. 652], [AJL+96, p. 10], [BJH+99, p. 445, 447-450], [Hen-ol]). During non-stationary strand deposition, the feed velocity of the plastic filament can be constant. If the delivered flow of plastic melt is not constant, the strand deposition is regarded as non-stationary.

If the delivered flow of plastic melt is not constant, the strand deposition is regarded as non-stationary. Stationary strand deposition is characterized by the attained constant target velocity of the translational motion and the delivery of plastic melt with a constant volume flow at constant feed velocity of the plastic filament.472Effects observed on the partThrough the influencing factors acting on the FDM process, part-specific structures are formed. Some can be seen even on the deposition of an individual strand while others only occur when several strands are laid together to create a strand package. A few examples are given below.Starting and ending of the strand depositionFDM parts contain numerous seams. Seams occur at all points of contact of deposited plastic strands. With seams formed through the starting or ending (or both) of a strand extrusion (cf. [Geb13, p.

Seams occur at all points of contact of deposited plastic strands. With seams formed through the starting or ending (or both) of a strand extrusion (cf. [Geb13, p. 71]), additional process operations can be performed, depending on the data processing and the available machine technology. The various process operations on starting or ending the strand deposition serve, for example, to produce well-fused seams that are as clean as possible. Other seams come about through the laying of the strands next to each other and on top of each other, and are usually not influenced by additional process operations.Fig. 1 shows a seam created by the starting and ending of a strand deposition. In the deposition strategy, the strand deposition shown here is defined as the outer contour of a part. In Fig. 1, the outer face of the component is pointing upwards.

In the deposition strategy, the strand deposition shown here is defined as the outer contour of a part. In Fig. 1, the outer face of the component is pointing upwards. The direction of motion of the tool during strand deposition as well as the translational motion of the tool afterwards are shown in the diagram with arrows.Outer faceDirection of depositionDirection of depositionNotchEndmove1 mmInner faceFig. 1:Macrograph of a seam on the outer face (produced from ULTEM 9085 with a T16 nozzle on a Stratasys Fortus 400mc)The notch in the deposited strand that can be seen in Fig. 1 results from an end-move of the tool after depositing the strand.

1 results from an end-move of the tool after depositing the strand. With this motion, the nozzle is advanced over the already deposited plastic strand into the inside of the part in order to produce as clean a weld as possible on the outer face of the component.Acceleration processes in strand depositionWhen starting and ending a strand deposition or when changing direction during strand deposition, the tool, because of dynamic forces, must be accelerated and decelerated. In modern machines, the feed velocity of the filament is adjusted during non-stationary strand deposition to the translational motion velocity of the tool. The aim is to ensure that a strand of uniform thickness473is deposited. Through various disruptive factors, such as the material-specific Young’s modulus and machine-specific spring length of the material, this process is at present not very reliable. In Fig.

Through various disruptive factors, such as the material-specific Young’s modulus and machine-specific spring length of the material, this process is at present not very reliable. In Fig. 2 it can be seen that, in areas of non-stationary strand deposition, with acceleration and deceleration processes of the tool (cf. [TG15, p. 253], [HJ00, p. 1711]), non-uniform strand thicknesses are produced.0 . 5 8 m m0 . 5 5 m m0 . 5 2 m mDirection of depositionDirection of deposition0000....5 0 m m5 2 m m5 3 m m5 6 m m2 mmFig. 2:Acceleration processes during strand deposition (produced from ULTEM 9085 with a T16 nozzle on a Stratasys Fortus 400mc)Strand deposition around a radiusWhen depositing a strand around a radius, a compression or thickening occurs on the inside of the strand and a stretching of the material on the outside. There are also areas in which the travel distances of the tool overlap.

There are also areas in which the travel distances of the tool overlap. Here, more material is deposited than necessary as a result of overlapping of the travel motions (cf. [HJ00, p. 1711]). The smaller the radius around which the strand is deposited, the more marked this effect becomes. Fig. 3 shows the deposition of a strand around a 90° angle. The deviation from the planned strand deposition plus a thickening of the inner radius and a lack of material on the outer radius can be seen.Direction of0.53 mmdepositionDirection ofdouble deposition areadeposition0.67 mm0.53 mmFig. 3:Strand deposition around a radius (produced from ULTEM 9085 with a T16 nozzle on a Stratasys Fortus 400mc)474Layer gaps due to the strand deposition strategyAt present, depending on the particular layer, a constant width is used for the deposited strands while the track width is allocated to the strand width.

If, in any particular area of the part, the deposited tracks of the inner screen are aligned parallel to the contour line, it is impossible to produce just any desired wall thickness. Depending on the settings for the strand deposition strategy, there are two possibilities for mapping the relevant area of the part. One possibility is to map the area of the part with a discrete wall thickness corresponding to many times the track width, in which case the finished part may not have the intended wall thickness. The other possibility is to deposit the strand for the contour in such a way that the outside measurement of the wall of the part corresponds to the intended size and to fill the inner areas by laying as many strands as possible next to one another. As long as the inner dimension cannot be mapped as a whole-number multiple of the track width, gaps will occur in this area of the part (cf. [HRE+13, p. 385]). Fig.

As long as the inner dimension cannot be mapped as a whole-number multiple of the track width, gaps will occur in this area of the part (cf. [HRE+13, p. 385]). Fig. 4 illustrates this effect with the aid of a cuboid produced by FDM with edge lengths of 5*5*25mm.5 mm0.508 mm0.722 mm5 mmFig. 4:Top view of a cuboid to visualize the strand deposition strategy (produced from ULTEM 9085 with a T16 nozzle on a Stratasys Fortus 400mc)Twin formation and variable notch geometriesIn FDM parts, it can be seen that strands lying next to one another group together to form pairs of twins. This means that both strands grow together to a certain extent and partially cover the same space. As a consequence of this, the resultant twin no longer has the full double width of a strand. A possible explanation for this behavior is that the strands group together in the molten state to reduce surface energy. This theory is supported by the sintering tests carried out by [BLS+04, p. 174ff].

This theory is supported by the sintering tests carried out by [BLS+04, p. 174ff]. The strands are deposited side by side, strand by strand. The first strand is deposited and solidifies at its respective position. Subsequently, the next strand is laid next to it in the molten state. The latter lies so close to the first, already solidified strand that the melt, under the compulsion to reduce the free surface, can pull slightly in the direction of the deposited strand and can solidify there. With the third strand, however, the gap to the second strand is so large that it can solidify in its intended position. This process repeats itself in the course of the build process.To illustrate this effect, the part shown in Fig. 5 on the left was produced by the FDM process. The strand deposition strategy was selected so that all strands are aligned in the direction of the symmetry axis.

5 on the left was produced by the FDM process. The strand deposition strategy was selected so that all strands are aligned in the direction of the symmetry axis. The formation of twins can be seen irrespective of whether the strands lie exactly above one another or with a gap (cf. Fig. 5 right). It can also be easily seen in Fig. 5 that,475in the production of a strand package, depending on the strand deposition strategy and shape of the part, very different notch geometries are formed inside a part.1 mmFig. 5:Formation of twins and notch geometries in the strand package (produced from ULTEM 9085 with a T16 nozzle on a Stratasys Fortus 400mc)Influences from the machine (kinematics, tool and build environment)There are a variety of FDM machines on the market. In some cases they differ considerably in their concept, design and construction.

In some cases they differ considerably in their concept, design and construction. Without taking the machine-specific influences into account it is not possible to determine characteristic data for evaluating the processing properties of different materials. Basically, an FDM machine can be divided up into three sub-assemblies: the kinematics, the tool and the building environment. With the help of these criteria, the most important influences are explained below.Kinematics – machine vibrations and positioning accuracyIn the FDM process, the kinematics have the task of moving the tool in relation to the building platform or the partially manufactured part. The positioning, track and repetition accuracy as well as the possible travel speed of the tool have a major influence on the deposition of the strands and differ significantly depending on the machines [Geb13, p. 80], [TSG14, p. 194].

80], [TSG14, p. 194]. It is essential that the strands are positioned very accurately next to one another and on top of each other in order to ensure reproducible weld seams and thus attain the specified strength data in the final part. In particular, depending on the machine stiffness and damping properties, vibrations occur in connection with the highly dynamic movements of the tool. As a result, the positioning and track accuracy are reduced after a period of vibration. This can also be seen in the production result and is illustrated in Fig. 6. The deposited strand was measured in the diagram at equidistant intervals. Deposition of the strand around the 90° change of direction and the acceleration processes taking place induces vibration of the machine construction. Following the change of direction this can also be recognized in the strand deposition in the form of thickness fluctuations.

Following the change of direction this can also be recognized in the strand deposition in the form of thickness fluctuations. These effects on strand deposition and machine vibration gradually decline.476Direction of deposition0.690.700.670.710.660.720.620.68mmmmmmmmmmmmmmmm2 mmFig. 6:Macrograph of a strand deposited around an angle of 90° (produced from ULTEM 9085 with a T16 nozzle on a Stratasys Fortus 400mc)Depending on the machine concept and quality of the machine calibration, the deposition of melt strands cannot be accurately performed directly on the build platform in the first layers of a part. This can lead to the width-to-height ratios (b/h ratios) of the deposited melt strands not corresponding to the original settings for the process. This in turn influences the density of the component, the width of the weld seams and the strength data of the part. Two error patterns can occur.

This in turn influences the density of the component, the width of the weld seams and the strength data of the part. Two error patterns can occur. If the first layer is set too low through the machine, the strands will be deposited with an excessive b/h ratio. The melt strands are squeezed on and the resultant weld widths are larger than with a correctly set layer height. If the layer is set too high, the b/h ratio is lower than intended. The strands are deposited on the platform and not squeezed on. The resultant weld widths (bs1,1 and bs1,2 in Fig. 7) are smaller than with a correctly set layer height.To illustrate the problem, Fig. 7 shows individual strands deposited on top of each other. On the left is a strand deposited with a correct layer height and a b/h ratio of 2. Fig. 7 right shows a strand deposition, which, with the same plastic delivery (i.e. the cross-sectional areas A1 and A2 are identical), results from the layer height being too high.

Fig. 7 right shows a strand deposition, which, with the same plastic delivery (i.e. the cross-sectional areas A1 and A2 are identical), results from the layer height being too high. The b/h ratio and the weld seam width are much smaller.b2,1b2,2bs1,1bs1,2zA1A22 s1 shx/yhFig. 7:b to h ratio and weld seam width with identical strand cross section477Through the process-specific deposition of a high-viscosity plastic melt to produce a porous part, the production process nevertheless has self-regulating properties. For example, the faults described above average out over several layers. It can be seen in Fig. 8 that, despite the initially excessive layer height, the target layer thickness is reached after only a few layers. The strands of the described part are deposited alternately with 0° and 90° orientation.h4h3h2h1h1 > h2 > h3 > h41 mmFig.

The strands of the described part are deposited alternately with 0° and 90° orientation.h4h3h2h1h1 > h2 > h3 > h41 mmFig. 8:Macrograph of the first layers of the part (produced from ULTEM 9085 with a 0.4 mm nozzle on an FDM laboratory machine)Plastic filament extruderThe plastic filament extruder shown schematically in Fig. 9 constitutes the tool in the FDM process. It conveys the semi-finished product, i.e. the plastic filament, plasticizes it and forces the plastic melt through a nozzle onto the build platform or already produced areas of the part [TSG14, S. 192], [AJL+96, S. 5]. Basically, the plastic filament extruder consists of the feed unit and plasticizing unit. Depending on the machine, the construction of this tool varies considerably. There may be differences e.g.

Basically, the plastic filament extruder consists of the feed unit and plasticizing unit. Depending on the machine, the construction of this tool varies considerably. There may be differences e.g. in the concept for conveying the filament, the length and design of the plasticizing zone and the arrangement of the filament feed unit and the plasticizing unit.𝑀𝑀𝑀𝑀𝑓𝑓𝑙𝑙z𝑠𝑠𝑙𝑙𝑄𝑄̇ x/yFig. 9:FDM filament extruder478Because of the principle of the process, the plastic filament is stretched under pressure because the counterpressure created in the plasticizing unit counteracts the force of the filament feed. Depending on the construction of the plastic filament extruder, the length lf between the plasticizing unit and the feed unit can vary considerably. Depending on the material being used, the stiffness of the processed plastic can also vary (cf. [YHG+97, p. 692f]).

Depending on the material being used, the stiffness of the processed plastic can also vary (cf. [YHG+97, p. 692f]). In the area of non- stationary strand deposition, this leads to a material delivery that is difficult to coordinate from a material and machine point of view (cf. [HJ07, p. 651-652, 657-658]). If the compressive stress in the plastic filament is too high or if the length lf is too large or if the material stiffness too low, it is even possible that a buckling of the filament will occur. A differently designed plastic filament extruder will exert an effect especially when starting and ending the strand deposition or when changing the speed of the plastic delivery. Through the compressive stress present in the filament, the material will still be forced out of the nozzle even if the feed velocity is reduced to zero, because the plastic filament is under stress and the tension is released like a spring (cf. [AJL+96, p. 10].Fig.

[AJL+96, p. 10].Fig. 10 shows a component that has the seam in every layer at the same place and is always touched at the same xy position by a part being built next to it. As a result, the plastic emerging uncontrolled from the nozzle conglomerates on the side wall of the part and builds a kind of stairs.S e a m1.5 mmFig. 10:Building defects on the outer wall through uncontrolled delivery of materialOven / Building environmentIn an industrial environment, as is realized in facilities of the firm Stratasys (cf. [Str00]), the established procedure is to place the strands deposited by fused deposition modelling into a heated environment such as an oven. This has several advantages. On the one hand, it means the shrinkage of the individual strands can be reduced, and, on the other, the weld seam strength is increased.

This has several advantages. On the one hand, it means the shrinkage of the individual strands can be reduced, and, on the other, the weld seam strength is increased. Different machines have a different homogenous temperature distribution in the oven and produce a different amount of convection there.Determining the weld seam strengthThe FDM process can be regarded as a constantly repeating welding process. To be able to produce parts that can be subjected to stresses and loads, it is necessary for the strands deposited by the FDM process and fused to one another to enter into a firm bond.

To be able to produce parts that can be subjected to stresses and loads, it is necessary for the strands deposited by the FDM process and fused to one another to enter into a firm bond. The weld seam strength is therefore an important criterion for the suitability of a material for processing by fused deposition modeling.479Designing the test method and specimensTo evaluate and be able to compare the suitability of different materials for processing, it is necessary to design a test method and a test specimen that allow a representative weld that can be reproducibly produced by the FDM process. If at all possible, the weld seam strength should be dependent only on unambiguous process parameters and the material used. Because of the previously described effects, however, some areas of the part are formed that cannot be reproducibly produced by all machines and materials and can therefore not be compared with one another.

For this reason, the area to be evaluated in the specimen should not contain any of these influencing factors.To limit the testing of the weld seam strength to an FDM weld that can be easily produced and is also reproducible, the specimen should have a wall thickness of exactly one strand width. This eliminates all influences that occur in the strand package and also allows full visual examination. An exact description of the strand geometry through the production of simple sectional views, is also possible. In addition, care is taken to ensure that the area to be examined is located only in areas of stationary strand deposition, also that machine vibrations have already ceased and the area does not contain any radii and seams. Translational tool movements with motions that do not deliver any plastic are selected in such a way that the nozzle is not moved directly above already deposited strands and thus no additional energy is introduced into the part e.g.

by heat radiation.Taking into account a design suitable for production for the FDM process, such as adequate part stiffness, the structure shown in Fig. 11 was designed and produced by the FDM process. The first layers of the structure are not in the testing area so as to avoid the influence of any inaccurately set initial layer height. Interruptions to the structure of the first layers make it possible to gradually build up relatively high part stiffness with short strand depositions. As a result, materials with high shrinkage in the FDM process can also be used to produce this structure.Fig. 11:Tensile bar in the Z direction480Tensile test specimens of Type 1B to [DIN527-3] were subsequently cut from the structure produced by the FDM process (cf. Fig. 11). The tensile test was carried out on a tensile testing machine according to the specifications in [DIN527-3].

Fig. 11). The tensile test was carried out on a tensile testing machine according to the specifications in [DIN527-3]. In order to unambiguously describe the geometry, it is also possible, in addition to the outer dimensions of the specimen, to produce a sectional diagram (cf. Fig. 12) to show the strand geometry.Material comparison PA6With the help of the developed test method, two types of PA6 were then compared with each other. They differ only in their viscosity. The type described below as PA6 NV has a low viscosity while the PA6 HV has a high viscosity. Tests were carried out with different process temperature pairs using otherwise constant process conditions in the area of stationary strand deposition (see Table 1).

Tests were carried out with different process temperature pairs using otherwise constant process conditions in the area of stationary strand deposition (see Table 1). The filament is dried before FDM processing.Table 1: Build parametersParameterFDM laboratory machineNozzle temperature [°C]250, 300, 350Oven temperature [°C]25, 75, 125Nozzle diameter [mm]0.4Layer height [mm]0.25Strand width set for the data processing [mm]0.5Speed in the area of stationary strand deposition [mm/s]50For every tested pair, a photomicrograph was produced to determine the strand geometry. Fig. 12 shows as an example the photomicrograph for the material PA6 NV at a nozzle temperature of 300 °C and an oven temperature of 75 °C.Width of the strandWidth oft he weld0.25 mmzFig. 12: Width of the weld and the strand, nozzle temp. 300 °C, oven temp. 75 °C, PA6 NVTo compare the materials with each other, a weld width factor is determined for all processpoints.

12: Width of the weld and the strand, nozzle temp. 300 °C, oven temp. 75 °C, PA6 NVTo compare the materials with each other, a weld width factor is determined for all processpoints. This is defined according to Equation 1:(1)Weld width factor =Width of the weld Width of the strand481Evaluation of the viscosity on the weld seam strengthFig. 13 shows the weld width factors for all parameter pairs for the materials PA6 NV (left) and PA 6 HV (right). The non-plotted test points with the material PA6 HV and a nozzle temperature of 250 °C could not be produced. The material has such a high viscosity that reliable melt delivery is not possible.

The non-plotted test points with the material PA6 HV and a nozzle temperature of 250 °C could not be produced. The material has such a high viscosity that reliable melt delivery is not possible. One striking aspect is that the weld width factor with the PA6 NV is dependent on the process temperatures, but that this effect was not observed with the PA6 HV.Weld width factors of PA 6 NVWeld width factors of PA 6 HV110,80,80,60,60,40,40,20,200Oven: 25°COven: 75°COven: 125°COven: 25°COven: 75°COven: 125°CNozzle: 250°CNozzle: 300°CNozzle: 350°CNozzle: 250°CNozzle: 300°CNozzle: 350°CFig. 13: Weld width factors of PA6 NV and PA 6 HVTo compare the welding seam quality of various materials, the tensile strength is shown in Fig. 14 related to the width of the weld (cf. Fig. 12). In this view, fracture-mechanical effects such as notch factors are neglected.

14 related to the width of the weld (cf. Fig. 12). In this view, fracture-mechanical effects such as notch factors are neglected. With the test points that have not been plotted, production or testing of the specimens was not possible.Tensile strengths in relation to the width of the weld for PA 6 NVTensile strengths in relation to the width of the weld for PA 6 HV5050404030302020101000Oven: 25°COven: 75°COven: 125°COven: 25°COven: 75°COven: 125°CNozzle: 250°CNozzle: 300°CNozzle: 350°CNozzle: 250°CNozzle: 300°CNozzle: 350°CFig. 14:Tensile strengths of PA6 NV and PA 6 HVSuitable process temperatures can be selected with the help of the building results and the measured tensile strengths. Under the condition that the process must be reliable and reproducible, only a nozzle temperature of 300 °C proved suitable.

Under the condition that the process must be reliable and reproducible, only a nozzle temperature of 300 °C proved suitable. At a nozzle temperature of 250 °C, there is a risk that the melt delivery is halted because the material viscosity is too high or it is blocked by482non-molten material. As long as a reliable melt delivery is possible, the attainable weld seam strengths are also low because of the low energy input. At a nozzle temperature of 350 °C, the material already becomes thermally damaged, causing production defects. The results of the tensile tests at a 350 °C nozzle temperature therefore have a high degree of scatter. An oven temperature of 125 °C is not suitable for the production of parts as the material turns yellow with time at this temperature.

An oven temperature of 125 °C is not suitable for the production of parts as the material turns yellow with time at this temperature. Because an oven temperature of 75 °C provides for the introduction of more energy into the welding seam and the material does not become thermally damaged or discolored at this temperature, a nozzle temperature of 300 °C and an oven temperature of 75 °C have emerged as optimum.With the help of the base material strengths determined on injection-molded specimens (see Table 2) and the performed tensile tests, a weld seam factor can be given. This relates the strength of the welding seam in the FDM process to the strength of the base material. Overall, it can be said that the PA6 NV material has a better weld seam strength than PA6 HV and also has a better weld seam factor.

Overall, it can be said that the PA6 NV material has a better weld seam strength than PA6 HV and also has a better weld seam factor. The tests thus show that, under the aspects of high weld seam strength, preference should be given to a low-viscosity PA6 type.Table 2: Weld seam factorsMaterialBase material strength [MPa]FDM weld strength [MPa]Weld seam factorPA 6 NV76.139.690.52PA 6 HV75.231.700.42ConclusionTaking account of the machine and process-related influences affecting the FDM process and the effects on the final product, a test method and a test specimen were produced. Through the specific design of the specimen and the simple and virtually complete documentation of the specimen geometry by producing a sectional view, a virtually machine-independent comparison of different materials is possible.

nstruction by Contour Crafting using Sulfur Concrete with Planetary ApplicationsBehrokh Khoshnevis, Xiao Yuan, Behnam Zahiri, Bin XiaDepartment of Industrial and Systems Engineering, University of Southern California, California, CA 90089AbstractThis paper reports on the experiments with the Contour Crafting Automated Construction process using sulfur concrete as the choice of construction material. Sulfur concrete has numerous terrestrial applications and is potentially an ideal construction material for planetary construction. On Mars, sulfur can be found in abundance and the range of temperature variation on the planet is within the safe zone for the structures to be built and survive over reasonable length of time with sulfur concrete. Several experiments have been performed at centimeter and meter scales. A FEA simulation model for the behavior of sulfur concrete based structures has been developed.

Several experiments have been performed at centimeter and meter scales. A FEA simulation model for the behavior of sulfur concrete based structures has been developed. Experimental results were compared with the results of simulation.Key words: additive manufacturing, Contour Crafting, sulfur concrete, planetary.IntroductionThe Sulfur Concrete Contour Crafting (SCCC) is a layer-based 3D printing technology which usessulfur concrete. Additive manufacturing has the advantage of having a low-cost, high strength, high acid/salt resistance and durable structure at high speed [1]. The SCCC is equipped with a 6-axis robot and a universal sulfur concrete extruder which can automatically collect local construction material and build on complex terrain. This technology has the potential to construct in outer spaces and in environments such as Mars or Moon, since both are abundant in sulfur.

This technology has the potential to construct in outer spaces and in environments such as Mars or Moon, since both are abundant in sulfur. [2]In the SCCC process the sulfur concrete composite materials (elemental sulfur, sulfur modifier, coarseaggregate and fine aggregate) are pre-melted and mixed at 150C, and are kept in the reservoir for one hour until the elemental sulfur are modified [1]. Then the well-prepared sulfur concrete is deposited by a special mixer/extruder fixed on a 6-axis robot. Sulfur concrete is deposited on the previous layer to the calculated path and cures in less than 10 minutes. After 24 hours, the construction made by sulfur concrete reaches its highest strength.BackgroundIn an earlier part of this NASA supported project a new mini-size auger with piezo vibration wasdeveloped to investigate the possibility of reliably extruding sulfur concrete.

As it is shown in Figure 1, dry sulfur and Martian regolith simulant in the funnel were pushed downward through a nozzle by a rotating auger while sulfur was being melted by electric heaters. The nozzle head formed the shape of the exiting sulfur concrete and built smooth surfaces and edges. To increase the bonding strength of the layer, a feature was added to the nozzle to produce an interlocking profile for successive layers [3].To automatically counteract the bridging effect during the extruding process a vertical vibration wasinduced by a piezo transducer to the auger and a horizontal vibration piezo was added to the end of the extruder shown in Figure 1. These two piezo actuators partially alleviated friction during the extrusion process.Figure 1: Schematic of Mini-size auger extruderTo improve the efficiency of heating, an experiment was carried out by pre-heating mixed sulfur and Martian regolith simulant in water at 98℃ for half an hour.

This dramatically improved the melting efficiency in the extrusion process as shown in Figure 2. In comparison to heating process in room temperature from melting temperature of sulfur at 115 C, pre-heating method further reduces the extrusion friction and improves the final surface of extruded Martian sulfur concrete. Of course use of water on Moon and Mars is problematic and other approaches such as the use of microwave for preheating may be considered. Figure 3 shows the Martian regolith dome built by this process.[3]Figure 2: Extruded sample with smooth surfaceFigure 3: Martian regolith domeAlthough the surface quality of the extruded layer has been optimized, the rotation torque has been excessive. Also, auger blade gets damaged when it is being used for 50 hours at a low temperature, especially on the upper part of the nozzle.

Also, auger blade gets damaged when it is being used for 50 hours at a low temperature, especially on the upper part of the nozzle. As shown in Figure 4, while the auger blade initially had a width of 3mm, it became 0.3mm due to abrasion caused by hard aggregate in regolith stimulant. Silicon dioxide makes half of the content of the regolith stimulants, having durability measuring between 4500~9500 MPa.Figure 4: Auger worn-out during experimentThe new approachTo ensure durability, stability and reliability of sulfur concrete extruder, a new universal concrete extruder was conceived and examined. As shown in Figure 5, the new mixer/extruder combination mechanism was devised in which the entire mixing and extrusion chambers were precisely heated. In contrast with the previous auger extruder, the stage-wise mechanisms on the upper sections move the material downward while a special extruder at the end of the nozzle provides the main extrusion force.

In addition, the mixture is completely melted before entering the nozzle and as such its friction with the walls of the nozzle is far less than the mixture at the ambient temperature or even pre-heated mixture. Besides these two alterations, an aluminum extender end which also acts as heat sink is added to the outlet of the nozzle to lower the temperature of the exiting material quickly. Fuzzy logic control is employed to control the DC motor that mixes the material and moves it downward through the nozzle. Although this control system has proven to be more consistent and accurate, the volume of extrusion tube limits the maximum length of extrusion.Figure 5: The mixer/extruder assemblyA continuous process for feeding the nozzle assembly was developed on the basis of the above approach. Figure 6 shows the new fabrication process concept where raw aggregate and dry sulfur are transported into a hopper installed on top of the nozzle. The hopper heats and mixes the material at 150 C.

Figure 6 shows the new fabrication process concept where raw aggregate and dry sulfur are transported into a hopper installed on top of the nozzle. The hopper heats and mixes the material at 150 C. As the sulfur concrete mixture in the reservoir gets consumed, the robot moves the extrusion system along with its reservoir under the silo outlet at the refill station.Raw material preparationMelting sulfur •Heating aggregateRefillingPositioningRobotExtrudingExtruder/RobotFigure 6: The construction processFor the construction of full-scale structures, a large industrial robot is being used in the Contour Crafting Laboratory. The robot (KUKA brand) consists of six rotational actuators which provide six degrees of freedom. One of the most important advantages of using such robot over a gantry robot is that for a given size of structure the gantry has to be designed in such a way that it can contain the structure within its work envelope.

However,a joint robot such as the KUKA robot used in this experiment can be mounted on top of a rover to eliminate the size restriction.Figure 7: Sulfur concrete contour crafting robotFigure 8: Three layers extrusion sampleBuilding 3D dome structures requires following more complicated pathways than straight or circular lines. For a stationary robot (not mounted on a rover) the position of starting layer in the global coordinate system must be selected in such a way that the whole dome structure falls completely within the robot workspace. Different shapes and geometries can be selected for the dome-like structures. One of the candidates for the 3D dome structure is depicted in Figure 9. For this structure inclined layers are built consecutively such that each layer has a slight angle with respect to the horizontal plane.Figure 9: Tool path coordinates for dome-like structureOther patterns to achieve dome-like structures are also being considered.

For instance, patterns which inherit geometric advantages of domes and cones can be used to build optimum structures by horizontal, rather than inclined, layering method. We plan to investigate relevant merits of different geometric structures in order to determine the optimum shape such that each layer has sufficient support and the whole structure shows robust stability and strength.Sulfur concrete extrusion studyWorkability is the most important property of molten sulfur concrete, which not only affects the transportation and extrusion but also the strength and surface quality of the final product. Bartos [4] explained workability with several fundamental characteristics such as viscosity, mobility, internal friction, pumpability, segregation, bleeding, formability and finishability.In Contour Crafting with sulfur concrete, mobility and pumpability of molten sulfur concrete affect the extrudability of the mix.

While, the formability and finishability affect the strength and surface quality of extruded strucutres. Unfortunately, the molten sulfur concrete with high mobility and pumpability always haspoor formability and finished quality and this is the difficult contradiction in this project. So, exploring the appropriate sulfur concrete for the two sets of contradicting properties and developing the effective nozzle for forming sulfur concrete are the two main tasks of this project [4].The performance of the extrusion process varies considerably with different proportions of sulfur in sulfur concrete. When sulfur proportion in the mix is high the sulfur concrete has low viscosity, which makes it easy to extrude. However, low viscosity also causes the problem of having high slump and leakage.

On the other hand, low proportion of sulfur lowers the slump, but a low viscosity mixture is hard to extrude as it always sticks to extrusion barrel and internal mechanisms.To explore the significant factors that may influence the extrusion process a 2k factorial experiment was designed based on temperature at different stages, proportion of sulfur and extrusion/moving speed as shown in Table 1.Table 1: A 2k factorial design on extruding process Level - 30 wt% 140 C 150 C 3 kg/min 3 kg/minFactors Sulfur Proportion Sulfur Concrete Temperature Extruder Temperature Extrusion Rate Extrusion RateLevel + 35 wt% 150 C 160 C 5 kg/min 5 kg/minWhen the sulfur concrete does not stick to the inner wall of the barrel and with the force of gravity flows smoothly to the lower sections of the nozzle, the experiment shows that the extruder temperature and extrusion rate have little effect on the extrusion process.(a) Sample built in 140C(b) Sample built in 150CFigure 10: Extrusion process with different sulfur cement temperatureIn the experiment it is obvious that the mixture at 150C has less viscosity and better flow rate.

Some extrusion cross-sections are observed under optical microscope, as shown in Figures 10 and 11. The porous structure is more equally distributed in the sample built at 150 C as shown in (b), while most of porous structure is shown in (a) as sample built in 140C. It is deduced that sulfur in mixture at 150C took longer for the curing process and sank into the lower region of parts by the force of gravity. The uneven sulfur content between upper and lower parts may considerably reduce the strength of multi-layer structures.(a) Sample built in 140C(b) Sample built in 150CFigure 11: Extrusion process with different sulfur cement temperatureSulfur plays the role of binder in solidified sulfur concrete and the role of lubricant during the extrusion process. By increasing sulfur proportion the viscosity of sulfur cement decreases considerably, therefore, the liquidity increases.

By increasing sulfur proportion the viscosity of sulfur cement decreases considerably, therefore, the liquidity increases. Compared to the 30 wt% sulfur concrete samples, the 35 wt% sulfur concrete is saturated with sulfur and hence the porosity level is not significant. However, the plasticity of 35 wt% sulfur cement in the molten state is less than that of 30 wt% sulfur concrete. This may explain the reason behind the round shapes of both edges and top surface of the sample on the right in Figure 12. In such cases it is especially necessary to add side trowels behind the nozzle to keep the extruded surface smooth.(a) Sample with 30wt% sulfur(b) Sample with 35wt% sulfurFigure 12: Extrusion process with different mixture temperature(a) Sample with 30wt% sulfur(b) Sample with 35wt% sulfurFigure 13: Extrusion process with different mixture temperatureFrom Figure 13 it is clear that the microscopic characteristics of the two samples are different.

The size of microscopic holes in the two samples differs greatly. Porosity directly impacts the strength of the structures made with these materials. By choosing the appropriate weight ratio for sulfur in the mixture, the porosity can be decreased.Finite Element AnalysisA finite element analysis for the deformation behavior of sulfur concrete immediately after extrusion was conducted by using the finite element code. The details of analysis are given below.a) Description of constitutive model and selection of material parameters Mohr Coulomb Plasticity model, provided within the material library of ABAQUS, was selected for the description of the rheological behavior of uncured sulfur concrete right after extrusion. According to Mohr- Coulomb model, failure in the element begins when the shear stress in a material reaches a threshold that depends linearly on the normal stress in the same plane.

According to Mohr- Coulomb model, failure in the element begins when the shear stress in a material reaches a threshold that depends linearly on the normal stress in the same plane. By plotting Mohr's circle for states of stress at failure in the plane of the maximum and minimum principal stresses, the failure line is defined as the best straight line that touches these Mohr's circles as shown in Figure 13 [5].Figure 13: Mohr-Coulomb failure model [5]According to figure 13, the Mohr-Coulomb model is defined by [5]𝜏 = 𝑐 − 𝜎 tan 𝜙(1)Where 𝜎 would have a positive sign in tension and a negative sign in compression. In Equation (1) 𝜙 is the friction angle and 𝑐 is the cohesion of the material.

With some manipulation in Equation (1), the Mohr- Coulomb model can be written as𝑠 + 𝜎𝑚 sin 𝜙 − 𝑐 cos 𝜙 = 0(2)This model can also be written in the general form as:𝐹 = 𝑅𝑚𝑐𝑞 − 𝑝 tan 𝜙 − 𝑐 = 0(3)Where𝑅𝑚𝑐(Θ, 𝜙) =1√3 cos 𝜙sin (Θ +𝜋3) +13cos (Θ +𝜋3) tan 𝜙(4)The essential components of the Mohr-Coulomb model are described below:Friction angle: In Equation (1-4), 𝜙 is the friction angle which is the slope of the Mohr-Coulomb yield surface in the 𝑝_𝑅𝑚𝑐𝑞 stress plane (see Figure 14), which is commonly referred to as the friction angle of the material and can depend on temperature and predefined field variables.Dilation angle measured in the 𝑝_𝑅𝑚𝑐𝑞 plane at high confining pressure and can depend on temperature and predefined field variables. Cohesion: In Equation (1-4) 𝑐 is the cohesion of the material.The geometry of the yield surface in the deviatoric plane is controlled by the friction angle 𝜙 as shown in Figure 14.

The range for friction angle variation is from 0 ≤ 𝜙 < 90. In the case of 𝜙 = 0, the Mohr-Coulomb model reduces to the pressure-independent Tresca model with a perfectly hexagonal deviatoric section. In the case of 𝜙 = 90 the Mohr-Coulomb model reduces to the “tension cut-off” Rankine model with a triangular deviatoric section and 𝑅𝑚𝑐 = ∞ [5].Figure 14: Mohr-Coulomb yield surface in meridional and deviatoric planes [5]b) Geometry and boundary conditions The geometry of the specimen is shown in Figure 15. Only one half of the specimen could also be modeled for the analysis to get the results since the deformation analysis of the concrete has axisymmetric nature. The specimen was meshed using four-node axisymmetric elements.

The specimen was meshed using four-node axisymmetric elements. The surface beneath the specimen was modeled as a rigid body with friction interaction with the model.Figure 15: The geometry of the specimen and the initial mesh used to implement the finite element analysis.c) Numerical results and a comparison with the experimental data The relative geometrical deformation from the FEA model was obtained and compared to the experimental results (Figure 16, 17). By changing the viscosity in each experiment and using an iterative recursive curvefitting method, we are able to obtain an FEA model for uncured sulfur concrete which can predict the deformation of the extruded mixture as a function of viscosity.Figure 16: The Mohr-Coulomb based FEA model and the extruded sample 35%wtFigure 17: The Mohr-Coulomb based FEA model and the extruded sample 30%wtFigures 16 and 17 show the FEA models versus the experimental samples.

The essential parameters in the Mohr-Coulomb model is different for the two samples with different sulfur contents. The appropriate friction coefficients between the samples and the base were found by comparing the experimental results with the FEA results so that the friction coefficients were in the best agreement between the predicted and experimental deformation patterns.6. ConclusionExperimental samples was developed using a novel mixer/extrusion system. The mechanism was proven to be durable and stable after more than 500 hour’s work. The extrusion experiment showed that the temperature and proportion of composite influenced the final shape of extrusion. Higher extrusion temperature made the samples less porous. Less sulfur proportion improved the shape and surface of the extrudate but increased the porosity. Relative deformation of different samples with different proportions of ingredients were studied.

Less sulfur proportion improved the shape and surface of the extrudate but increased the porosity. Relative deformation of different samples with different proportions of ingredients were studied. The experimental results obtained were compared with those found from a FEA analysis. A constitutive model based on Mohr Coulomb Plasticity model was used for the FEA calculations. The simulations based on the FEA model show the potential for improving the deformations of sulfur concrete after extrusion.AcknowledgementThe research reported here has been supported by a Phase II NASA Innovative Advanced Concept (NIAC) program.Reference[1] A.-M. O. Mohamed and M. El-Gamal, Sulfur concrete for the construction industry: a sustainabledevelopment approach. J. Ross Publishing, 2010.[2] B. Khoshnevis, M. Bodiford, K. Burks, E. Ethridge, D. Tucker, W. Kim, H. Toutanji, and M. Fiske,“Lunar contour crafting—a novel technique for ISRU-based habitat development,” Autom. Constr., vol.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCONSTRUCTION OF METALLIC GLASS STRUCTURES BY LASER-FOIL- PRINTING TECHNOLOGYYiyu Shen, Yingqi Li, and Hai-Lung TsaiDepartment of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractMetallic glasses (MGs) have superior mechanical properties such as high tensile strength, hardness, and corrosion resistance, as compared to crystalline metals. Although newly developed MGs have significantly reduced critical cooling rates down to 10 K/s, products of MGs are still limited to simple geometries such as foils/plates or rods with thin section-thickness which is mainly caused by the decrease of thermal conductivities of the new MGs.

Recently, we developed a new Laser-foil-printing (LFP) additive manufacturing technology which welds foils, layer by layer, to construct desired 3D structures. With the LFP and Zr-based amorphous foils, 3D, large amorphous structures with complex geometry have been successfully manufactured. To better understand the evolution of crystalline phase, we integrate the finite element based heat transfer model and classic nucleation theory (CNT) based crystal nucleation/growth model. The model was used to demonstrate the evolution of crystal phase as a function of time during laser welding at different locations including the fusion zone (FZ) and heat-affected zone (HAZ). The model is also compared favorably with the experiment results. The reported susceptibility to crystallization in HAZ were discussed and explained.IntroductionMetallic glasses (MGs) or amorphous alloys can be produced by rapid quenching and solidification of molten metals to bypass crystallization.

Atoms in amorphous metals exist in a random/disordered manner much like atoms in the liquid phase. Compared to traditional crystalline metals, MGs have superior properties such as high tensile strength, hardness, wear resistance, and corrosion resistance [1−5]. These properties are inherited from randomness/disorder of atoms in the MG which is different from crystallized metals in which periodic arrays of atoms, called crystal lattices, are arranged in an organized, structured manner. In a crystalline metal, however, mainly due to multi-directional slow cooling and solidification, defects/dislocations/segregations always exist between grains/crystals (grain boundaries or interfaces) which are the “weak spots” and could initiate “cracks” when subjecting to external loading, propagating and eventually leading to the failure of the entire structure.

Hence, tensile strength, for example, for a “normal” crystalline metal can only be a fraction of otherwise “perfect” crystal lattices of the same metal composition.The main challenge of producing large metallic glass part/structure is that a certain high cooling rate must be reached otherwise the alloy would been crystallized before it cools down to room temperature. This is named as the critical cooling rate (Rc). Once crystallized, the unique mechanical properties could be lost. Limited by the manufacturing method, this critical cooling rate also limits the size of MG part to be produced. One example is the casting method, which is currently the most widely used method for produce MG parts in this area. During casting the755material who is away from the surface cools less fast compare to that close to the outer surface. Thus, a defined critical cooling rate is equivalent of an upper limit of casting thickness.

Thus, a defined critical cooling rate is equivalent of an upper limit of casting thickness. Additive manufacturing, however, build up the designed part layer by layer. The success application of AM technology on MG structures will separate the cooling process into every individual layer. As a result, once the feasibility of AM for metallic glass is proved, the limitation of dimension in manufacturing metallic glass product can be broken through. Unfortunately, to date no successful result has been published. Given that the outstanding mechanical properties and high application potential of MG, fundamental research is required to address the significant technological challenges and bridge the capability in AM of MG structures.In laser material processing, as a large amount of energy is tightly focused onto a very small area, the heating rate (Rh) and Rc of the processed part are very high (104 ~ 105 K/s) [6].

This makes laser-based AM technologies naturally a promising method in building large MG parts [7]. Recently, Sun [8] studied the microstructure evolution of Vitreloy 106a MGs processed by LENS technology. He pointed out that the crystallization phenomenon in the heat-effected zone (HAZ) took place even at a Rc of 104 K/s due to complex thermal history. Hence, preventing crystallization in the HAZ is the main challenge rather than in the melting zone. Zhang et al. [9] have found a similar phenomenon in powder-based AM process. Therefore, it is vital important to understand the crystallization mechanism behind the laser-MG interaction. This is different from the posted researches on regular crystallization of MGs[11-13].In this paper, we first demonstrate the use of developed laser foil printing (LFP) technology to weld amorphous foils, layer-by-layer, via laser to construct 3D MG parts. Then, the evolution of crystallization phase during the laser welding on Zr- based MG is studied.

Then, the evolution of crystallization phase during the laser welding on Zr- based MG is studied. A finite element based heat transfer model is applied to obtain the temperature history in the HAZ. After that, classic nucleation theory based nucleation/growth model is applied to calculate the evolution of crystalline phase.The laser foil printing (LFP) technologyDetails of the LFP technology can be found in Ref. [14]. However, in order to better demonstrate how a part is manufactured, which will affect the formation of crystalline, and also to discuss the effect of process parameters, the method, system, and procedure for the LFP technology are briefly described below. A schematic of the LFP process is given in Fig. 1. The LFP process is mainly realized by two alternative processes: foil-welding by a continuous wave (CW) fiber laser and foil-cutting by a pulsed UV laser.

1. The LFP process is mainly realized by two alternative processes: foil-welding by a continuous wave (CW) fiber laser and foil-cutting by a pulsed UV laser. A fresh foil is first loaded on the substrate and then fixed by spot-welding, the program controls the scanner to scan a specified pattern path of the CW laser beam on the foil to weld the foil to the substrate. Depending upon the desired Rc for different amorphous alloys, welding can be accomplished via line-welding or a matrix of spot-welding. The welding path/pattern should be designed (which depends on the geometry of the part) to achieve symmetrical welding in order to minimize possible thermal stress/part distortion, and at the maximum separations in time and in space between two consecutive weldings to avoid heat accumulation and achieve high cool rates. For example, for a spot-welding at location A, the next spot-welding at location B should be as far as possible from A.

For example, for a spot-welding at location A, the next spot-welding at location B should be as far as possible from A. After the welding process is done, an X-Y motion stage shifts the part to the location under the UV laser beam. Then, the program controls the X-Y motion stage and the UV laser to cut the welded foil along the outer and/or inner756circumferences of the part. After the laser cutting is completed, the redundant foil is removed and another fresh foil is placed and fixed on the top of the sample. The aforementioned entire process can be fully automatic. This procedure is repeated until the whole desired 3D part is built. Note, in the two alternative processes, laser welding can be done ahead of laser cutting or vice versa.Fig. 1. Schematic of the LFP process and the LFP system setups at the Laser-Based Manufacturing Lab, Missouri S&T.Figure 1 also shows the LFP setup developed at the Laser-Based Manufacturing Lab, Missouri S&T.

The welding laser is an IPG YLR-1000 CW (continuous wave) fiber laser with a wavelength of 1070 nm and a maximum output power of 1000 W. The cutting laser is a Coherent AVIA-355X laser with 355 nm UV wavelength, Q-switched, 30 ns pulse duration and 10 W maximum power. The UV laser, not shown in the figure, is located in another room. The welding laser beam passes through an optical fiber, enters a collimator to become an expanded parallel beam, then enters the scanner (hurrySCAN25, SCANLAB), and finally is focused on the foil through an F-theta lens with an effective focal length of 33 cm. The cutting UV laser beam passes through several turning mirrors and is focused by a focusing lens with a 10-cm focal length. A cooling chamber with a heat exchanger plate at the bottom by passing through liquid nitrogen or liquid argon is designed757to achieve rapid cooling while the amorphous structure is being manufactured.

The collimator and scanner for the fiber laser, and the focusing lens for the UV laser are mounted on a Z (vertical) motion stage. The foil is first positioned and fixed onto the substrate via spot-welding using a mold plate with predetermined hole-pattern, and the entire structure is placed on an X-Y (horizontal) motion stage (Aerotech). The two lasers, scanner and three-axis motion stage are integrated through control programs to a computer.Fig. 2. As-fabricated amorphous 3D parts by LFP.Figure 2 shows a few 3D MG parts printed by LFP from our previous publication [24]. Fig. 2(a) is a side view of a “metal fusilli pasta” part created by a continuous rotation of an ellipse pattern. Fig. 2(b) shows a 3-mm-thick amorphous “S&T” logo of the Missouri University of Science and Technology. A 5-stage 3D table, includes 4-square stages and a semi-spherical cover on the top is shown in Fig. 2(c). The overall height of the part was measured to be 11.3 mm. Fig.

A 5-stage 3D table, includes 4-square stages and a semi-spherical cover on the top is shown in Fig. 2(c). The overall height of the part was measured to be 11.3 mm. Fig. 2(d) is a slipped cylinder with overall displacement of 13 mm. All 4 samples were examined to be fully amorphous by XRD.Experiment and ResultsCast Zr-based MG alloy, LM105 (Zr52.5Ti5Al10Ni14.6Cu17.9 (at. %)) was used in this study. The cast Zr-based MG alloy, as samples, with a thickness of 0.75 mm, were machined into several specimen, each with 10 mm in length and width. All sample surfaces were polished with 800-grit SiC paper right before the experiment. Laser welding is applied (bead on plate) on the prepared samples. The laser used in this study is a continuous-wave (CW) fiber laser (model YLR-1000, IPG Photonic Corp.) with output wavelength of 1070 μm. The laser beam diameter D on the focal plane is 430 μm.

The laser used in this study is a continuous-wave (CW) fiber laser (model YLR-1000, IPG Photonic Corp.) with output wavelength of 1070 μm. The laser beam diameter D on the focal plane is 430 μm. The motion of the laser beam is controlled by a scanner (hurrySCAN III, SCANLAB) with the maximum scan speed of 2000 mm/s at the focal plane. All experiments were conducted in an argon shielded chamber. To determine suitable processing parameters, the laser power (P) was set to 100 W, while the laser scan speed (v) was varied from 10 mm/s to 500 mm/s.758As the main work of this study is to melt the surface of an MG sample at a depth of less than 200 µm, the laser power was selected at a relatively low value to avoid possible formation of keyholes during welding.

The laser processed samples were cross-sectioned, polished and then etched by a solution of 100 ml H2O, 2 ml H2O2 and 0.1 ml HF for 10 seconds and then examined by an optical microscope (OM) and scanning electron microscope (SEM, Helios S600).Fig. 3. Top and cross-sectional views of laser treated MG at a laser scan speed of 10 mm/s.Figures 3 shows the top and cross-sectional views of the laser treated surfaces at laser power P = 100 W and scan speed v = 10 mm/s. As shown in Fig. 3(a), a 700 μm wide FZ was created on the top surface with ~100 μm wide HAZ on each of its two sides. From the corresponding cross section image in Fig. 3(b), along the vertical centerline the depth of FZ was measured as 235 μm. The thickness of the HAZ reaches a maximum of approximately 140 μm directly below the FZ. The micro-crystalline (MC) phases can be clearly seen pre-existed embedded in the amorphous matrix shown as the dark-spots in the image.

The micro-crystalline (MC) phases can be clearly seen pre-existed embedded in the amorphous matrix shown as the dark-spots in the image. A back scattering SEM image (Helios, ETD detector, BSE mode) for an enlarged view of typical MC crystals is given in inset (bottom left of Fig. 3(b)). For the as cast LM105, the volume fraction (VF) of MC crystalline was measured as 2.14%. The volume fraction of crystalline phase were estimated from the OM image of the etched cross section surface with ImageJ software [20−23]. Two selected areas in HAZ in Fig. 3(b) are enlarged and shown in Figs. 3(c) and 1(d). In HAZ there are clearly three distinct regions, as shown in Fig. 3(c); region I, as marked, is next to the FZ. An obvious difference in color change between HAZ and FZ can be seen after etching. While in region I there are no obvious MC particles observed and its thickness is about 16.0 μm. In region II, however, spherical-like MC crystals are observed with an average diameter of 4.34 μm.

While in region I there are no obvious MC particles observed and its thickness is about 16.0 μm. In region II, however, spherical-like MC crystals are observed with an average diameter of 4.34 μm. Region II is about 64.4 μm thick. In region III, sub-size crystalline phases are clearly seen. The average diameter of the phase is about 1.34 μm and the thickness of region III is approximately 53.8 μm. (Note the thickness of HAZ slightly dependents on the direction. In order to keep consistent with further discussion, we choose the downward direction as shown in Fig.3 (b).)The distinct appearances of the three regions in HAZ is believed to be caused by different thermal histories experienced in each region during the laser processing. Materials in region I are most close to FZ, so they experienced temperatures close to melting. Based on Ref.759[17], the melting temperature of LM105 is 1085 K.

Materials in region I are most close to FZ, so they experienced temperatures close to melting. Based on Ref.759[17], the melting temperature of LM105 is 1085 K. Thus, region I may has reached to the temperature point beyond the maximum crystalline grow rate, resulting a relatively small crystal size. In region II, materials were rapidly heated to around 1000 K at which the crystal growth rate is about the maximum and, as a result, the size of crystalline phase is larger than those in other two regions. In region III, the materials reach the temperature at which the nucleation rate is high and near its maximum. Hence, dense but smaller crystalline phases are expected. To understand the aforementioned phenomena, detailed analysis and explanation are postponed to next section with the help of mathematical modeling.In Fig. 3(d), a cluster of MC crystal is observed near the top-right corner of the FZ.

3(d), a cluster of MC crystal is observed near the top-right corner of the FZ. The size of the cluster is about 40 μm in diameter, which is significantly larger than the pre-existed MC crystal. We believe the formation of this MC crystal in FZ is caused by the follow reasons. Amorphous phase has a higher energy status as compared with crystalline phase (denoted as Gibbs free energy, ); hence, to be melted, the crystalline phase needs more thermal energy. Assuming the laser heat deposition rates are the same for crystalline phase and its nearby amorphous phase, the crystalline phase requires a longer time to get melted as compared to the nearby amorphous phase. In other words, crystalline phase may still stay in its solid status after nearby amorphous phase is melted. For a Gaussian-like beam profile, at the edge of the FZ (melt pool), the crystalline phase may not have enough laser energy to get fully melted as compared to its nearby amorphous phase or that in the center of the FZ.

The un-melted MC phase acts like a pre-existed crystal, and the size of the crystalline phase may grow and becomes larger during the cooling process.∆𝐺𝐺Fig. 4. Cross-section view of laser treated MG at laser scan speeds of 30 mm/s, 50 mm/s and 100 mm/s.Figure 4 shows laser treated cross-sections at v = 30, 50 and 100 mm/s, respectively, for Figs. 4(a), 4(b) and 4(c). The size of FZ decreases with the increase of scan speed, as expected. Fig. 4(d) is an enlarged image for v = 100 mm/s, Fig. 2(c). The size of HAZ also decreases with the increase of v. The HAZ is no longer observable in Figs. 4(c) and 4(d). Further increase of the laser scan speed causes a dramatic decrease of the size of FZ and thus, there is no need for further increase of laser scan speed beyond 100 mm/s.760The crystallization process during laser-MG interaction is more complex than that during casting process.

In the casting process, the temperature cools directly from above the melting temperature to below the glass transition temperature, which can be simplified to a value of critical cooling rate. While in the laser-MG interaction, the temperature first rises from room temperature to above melting and then followed by a rapid cooling back to below glass transition temperature. Thus the critical cooling rate may not be sufficiently applied in the laser-MG interaction case. To better understanding the crystallization evolution during laser-MG interaction process, we setup a coupled model combined with heat transfer and classic nucleation theory (CNT) in this study. The heat transfer model is used to estimate the temperature history at the selected point, while the CNT model is used to estimate the volume fraction evolution of crystalline phase in the amorphous matrix.The heat transfer model and the related assumptions and validation have been discussed in our previous paper [24].

The laser power is assumed to be a Gaussian distribution at the top surface (z-plane) which is expressed as:𝜕𝜕𝜕𝜕 𝜕𝜕𝜕𝜕 = 𝜖𝜖𝜖𝜖(𝑇𝑇44 − 𝑇𝑇∞) + ℎ(𝑇𝑇 − 𝑇𝑇∞)−𝑘𝑘(3) represents the laser power density, α is the surface absorptivity, P is the laser power, r is the laser beam radius and v is the laser scan speed. The x-, y-, x-coordinate system was established at the center of the top surface for the plate with the dimensions of 10 mm by 10 mm by 0.75 mm. The calculated domain was: x (−5 mm, 5 mm), y (−5 mm, 5 mm), and z (0 mm, −0.75 mm). The laser scanned from (x = −5 mm, y = 0 mm, z = 0 mm) toward the positive x direction.

The calculated domain was: x (−5 mm, 5 mm), y (−5 mm, 5 mm), and z (0 mm, −0.75 mm). The laser scanned from (x = −5 mm, y = 0 mm, z = 0 mm) toward the positive x direction. From the heat transfer model, the temperature history at any location can be calculated.−2[(𝑥𝑥−𝑣𝑣𝑣𝑣)2+𝑦𝑦2] 𝑟𝑟22∗𝑃𝑃 2where𝐼𝐼𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙 = 𝛼𝛼 ∙ 𝐼𝐼𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑒𝑒𝜋𝜋𝑙𝑙For the nucleation model, the classic nucleation rate writes [25]:∗∆𝐺𝐺 𝑘𝑘𝜕𝜕 )whereis the nucleation rate,is a constant,𝐼𝐼(𝑇𝑇) = 𝐴𝐴 ∙ 𝐷𝐷 ∙ exp (−Boltzmann’s constant, 𝐼𝐼 barrier of nucleation.denotes the absolute temperature and𝐴𝐴𝐷𝐷obeys the following relation:(4) is the is the thermodynamic potentialis the effective diffusivity,∗𝑘𝑘𝑇𝑇 ∆𝐺𝐺∆𝐺𝐺∗(5) is the difference of Gibbs free energy between the solid and liquid (amorphous) phase.

The temperature dependent effective diffusivity, , calculated from measured viscosity in Ref. [27] by Stork-Einstein relation [26]. The Gibbs free energy difference between undercooled liquid and corresponding crystal solid, , as a function of temperature given in Ref. [28] was used in the present study. The reported nose temperature of 860 K from Ref. [29] was used. By evaluating the maximum value of I × U3 which corresponds to the nose temperature in the TTT curve, can be determined as 0.076 J/m2 [30]. Then, the constant, A, can be determined by running Eq. (7) with N = 0 at the nose temperature and examining the time of crystallization which matches the TTT curve [26,29]. The constant A was found to be 2.5×1033. Finally can be determined by running Eq. (7) using the thermal history in manufacturing from the liquid phase to room temperature.

The constant A was found to be 2.5×1033. Finally can be determined by running Eq. (7) using the thermal history in manufacturing from the liquid phase to room temperature. However, as it is difficult to know the exact thermal history during manufacturing; instead, we used the evaluated VF of the sample (2.14%) and predicted the equivalent averaged cooling rate during manufacturing which is about 50 K/s. Thus, the pre-existed nucleus density = 1×1014 m-3. For a was determined by integrating the nucleation rate w.r.t. time, which is selected location, by plugging the temperature history T(t) obtained from the heat transfer model, Eq. (1), into Eqs. (4) and (6) to calculate , respectively, and, then, the VF of crystalline phase, , as a function of time was computed by Eq. (7). Hence, for the entire sample domain, the VF of crystalline phase can be obtained as a function of time and location.𝐷𝐷∆𝐺𝐺𝑁𝑁𝑁𝑁and𝐼𝐼(𝑡𝑡)𝑈𝑈(𝑡𝑡)𝑋𝑋(𝑡𝑡)) and crystalline Fig. 5.

(7). Hence, for the entire sample domain, the VF of crystalline phase can be obtained as a function of time and location.𝐷𝐷∆𝐺𝐺𝑁𝑁𝑁𝑁and𝐼𝐼(𝑡𝑡)𝑈𝑈(𝑡𝑡)𝑋𝑋(𝑡𝑡)) and crystalline Fig. 5. Solid lines: The calculated nucleation rate ( growth rate ( ) as a function of temperature. Dashed line: calculated temperature history of laser scan with P = 100 W and v = 10 mm/s at different locations. Curve (1): bottom of HAZ (region III in Fig. 1(d)), Curve (2): middle of HAZ (region II in Fig. 1(c)), and Curve (3): in FZ.𝐼𝐼(𝑇𝑇)𝑈𝑈(𝑇𝑇)762𝜖𝜖(7)as a function of temperature for LM105 alloy. It is seen the peak of nucleation rate occurs at the temperature of about 780 K. However, the maximum crystal growth rate occurs at about 1030 K, and then drops rapidly to 0 at the melting temperature Tm = 1085 K. Figure 5 also plots the calculated temperature histories for three points in the sample at laser power P = 100 W and scan speed v = 10 mm/s.

Figure 5 also plots the calculated temperature histories for three points in the sample at laser power P = 100 W and scan speed v = 10 mm/s. The three dashed curves represent the heating and cooling histories for the points at (x = 0 mm, y = 0 mm, z = −0.35 mm), (0, 0, −0.3), and (0, 0, −0.18 ), respectively. For Curve (1) representing the heating and cooling for point (0, 0, −0.35), which is located in HAZ but far away from FZ (region III in Fig. 1(c)), it is heated to about 810 K by laser in 0.1 s and then followed by rapid cooling to below 500 K in around 0.5 s. The curve (point) experiences much nucleation while touches only a little bit of crystal growth, which implies that there would be high dense with relatively small crystalline phase generated. For the point (0, 0, −0.3) as presented in Curve (2) of Fig. 3 (region II in Fig. 1(c)), it reaches about 920 K. This is the temperature that U is high and one would expect larger sizes of crystalline phase to be observed here.

3 (region II in Fig. 1(c)), it reaches about 920 K. This is the temperature that U is high and one would expect larger sizes of crystalline phase to be observed here. For Curve (3), which is for the point (0, 0, −0.18) located in FZ, the temperature exceeds the melting temperature in around 0.05 s. It is reasonable to assume that all crystalline phases/amorphous/nuclei will be melted while melting occurs (see also in Fig. 6). In the cooling process, the liquid alloy first experiences the temperature range in which U is very high in the temperature range, 920 K ~ 1080 K. However, as it was free of nuclei in the newly solidified alloy, the growth of crystalline phase is not significant. Then, the alloy experiences the temperature range where nucleation is high, but the growth is low, leading to a certain amount of nuclei in the alloy after cooling to the room temperature. The aforementioned discussions explain and consistent with the results obtained from experiments as shown in Fig.

The aforementioned discussions explain and consistent with the results obtained from experiments as shown in Fig. 3 at different locations in the sample.Figure 5 plots the calculatedand𝐼𝐼𝑈𝑈From Fig. 5, if an amorphous alloy is cooling down from its molten state, as there is no nucleation site even the cooling curve hit the crystal growth curve, U(t), the volume fraction of crystalline phase, X(t), is very minimum until it hit the nucleation curve I(t) at which nucleation sites start to be generated. However, at this time the crystal growth rate decreases to small and, hence, the X(t) is limited. The dependence of X(t) is not very critical for cooling from the molten state, particularly during the manufacturing process (e.g., mold casting). However, if a solid amorphous alloy is heated from room temperature by, for example, a laser, then the location and its heating/cooling rates would be very critical for the X(t).

However, if a solid amorphous alloy is heated from room temperature by, for example, a laser, then the location and its heating/cooling rates would be very critical for the X(t). If a point is located in the HAZ, it would experience twice of passing through the I(t) and U(t) curves during its heating and cooling processes and, hence, the accumulated X(t) is great. This explains the reports that why crystallization occurred in HAZ even the cooling rate is much higher than the critical cooling rate for amorphous phase [9−13]. Based on the aforementioned discussion, it is easy to understand that the pre-existed nucleus density, N, would have a significant effect on X(t) in HAZ.763Fig. 6. Calculated the evolution of volume fraction of crystalline in FZ.in FZ for the same condition shown in Fig. 3(b) with laser power P = 100 W and scan speed v = 10 mm/s. The location studied is (0, 0, −0.18), also the same as Curve (3) in Fig.

3(b) with laser power P = 100 W and scan speed v = 10 mm/s. The location studied is (0, 0, −0.18), also the same as Curve (3) in Fig. 5, which is in FZ (the calculated depth of fusion zone is 230 μm from the top surface). The dashed curve represents the temperature history from the heat transfer model. The “discontinue” solid curve in Fig. 4 represents the evolution of as a function of time. The VF of crystalline increases rapidly upon heating to about 38% at t = 0.192 s. Then, melting starts and clears the crystalline/nuclei at t = 0.192 to 0.232 s. Upon cooling and starting at t = 0.232 s, grows from 0 again and ends up with a near constant value of 2.6×10-6 at t = 0.4 s. The final 𝑋𝑋(𝑡𝑡) residual nucleus density in the sample, Nres, is calculated as 2.23×1013 m-3 which is only 22% as = 1×1014 m-3. Note, this number of residual nucleus compared to the initial nucleus density density is equivalent to the pre-existed nucleus density for the later thermal circles.

Note, this number of residual nucleus compared to the initial nucleus density density is equivalent to the pre-existed nucleus density for the later thermal circles. 𝑁𝑁Figure 6 shows the calculated𝑋𝑋(𝑡𝑡)𝑋𝑋(𝑡𝑡)Fig. 7. Calculated the evolution of volume fraction of crystalline in HAZ at different depths.at different locations in HAZ shown in Fig. 1(b) with P = 100 W and v = 10 mm/s. Note in the experiment, the depth of FZ was measuredFigure 7 plots the evolution of VF of crystalline𝑋𝑋(𝑡𝑡)764as 235 μm and followed by an about 140 μm thick of HAZ (also see Fig. 8). In the modeling, the depth of FZ was calculated as 230 μm. From Fig. 7, for a point in HAZ but just right below FZ (0, grows to 100% rapidly. Unlike what happened in FZ, melting, which is able to 0, −0.23), at any location in HAZ increases and then reset the crystalline percentage to 0, the value of reaches a steady-state constant value through the entire heating-cooling process.

The VF of crystalline phase ends up with 98.82%, 48.14%, 6.74 %, and 0.76 % at the depth of 310 μm, 330 μm, 350 μm and 370 μm, respectively. It is considered to be undetectable (0.76 % < 2 %) at the 370 μm depth for X(t). Thus, the calculated thickness of HAZ is around 135 μm (~365 µm – 230 µm), which is consistent with the experimental result, ~140 µm, as shown in Fig. 3(b).𝑋𝑋(𝑡𝑡)𝑋𝑋(𝑡𝑡)The curve at the depth of 230 μm from t = 0.21 to t = 0.23 s, as shown in the enlarged inset on left-top, indicates a “dent” in the crystallization rate curve between t = 0.21 s and 0.22 s. This “retardation of crystallization” is in fact expected [31] and can be well explained with the help of Fig. 5. From Fig. 5, it can be seen that for temperatures below (but close to) Tm (e.g., 1035 K to Tm = 1085 K), U(t), the crystal growth rate decreases rapidly from its maximum to almost 0.

5. From Fig. 5, it can be seen that for temperatures below (but close to) Tm (e.g., 1035 K to Tm = 1085 K), U(t), the crystal growth rate decreases rapidly from its maximum to almost 0. Hence, for a point in HAZ but next to FZ if its heating history is such that the maximum temperature is very close to the melting temperature (but not yet melted), the crystallization process freezes. The retardation of crystallization is also frequently observed from the well-known TTT curve of metallic glasses. As can be seen in the right inset image of Fig. 7, at temperatures just below Tm (the shaded area) the supercooled liquid can stay in its amorphous status for a long time because of the frozen of crystallization. Thus, to the position next to the fusion zone, the average crystalline size would expected to be smaller. Note, this phenomenon may be only observed under rapid heat/cooling to the sub melting temperature like in laser welding.

Note, this phenomenon may be only observed under rapid heat/cooling to the sub melting temperature like in laser welding. Under a relatively slow heating/cooling, fully crystallization would be reached already before the temperature gets high enough [32]. An example is given in Fig. 8 where the final VF of crystalline is plotted as a function of depth from the surface, as shown in the inset, for laser scan speeds of 10 and 30 mm/s. The region on the left where VF = 0 represents the fusion zone. For v = 10 mm/s, the VF of crystalline reaches 100% at the depth of 223 μm, which is the start of HAZ just below the FZ. With the increase of depth, the VF of crystalline first keeps steady (100%) and then starts to drop at the depth about 280 μm to below 2% at the depth of 346 μm. If the criteria of fully amorphous is set to the final VF of crystalline < 2%, the thickness of HAZ is measured as 123 μm (346 µm minus 223 µm) under laser scan speed v = 10 mm/s.

If the criteria of fully amorphous is set to the final VF of crystalline < 2%, the thickness of HAZ is measured as 123 μm (346 µm minus 223 µm) under laser scan speed v = 10 mm/s. While for scan speed v = 30 mm/s, the final VF of crystalline is found to be about 43% right below FZ at the depth about 175 μm. It is noticed that the final VF of crystalline increases to about 50% at the depth about 182 μm and then drops back to 0 for higher depths, resulting in a 72 μm thick of HAZ (247 µm minus 175 µm). The maximum VFs of crystalline at other laser scan speeds are given in Fig. 8. As shown in Fig. 8, when the laser scan speed is about 17 mm/s or lower, the maximum VF of crystalline reaches 100%.765Fig. 8. Calculated final volume fraction of crystalline as a function of depth at different laser scan speeds.Fig. 9. Calculated maximum crystalline VF at different laser scan speeds.766Fig. 10.

8. Calculated final volume fraction of crystalline as a function of depth at different laser scan speeds.Fig. 9. Calculated maximum crystalline VF at different laser scan speeds.766Fig. 10. Calculated and measured depth of fusion zone (FZ) and thickness of heat-affected zone (HAZ) as a function of laser scan speed at a fixed laser power (P =100 W).The depth of FZ and the thickness of HAZ from experiments and modeling are shown in Fig. 8. For a fixed laser power P = 100 W, the scan speed v increases from 10 mm/s to 100 mm/s with an increment of 10 mm/s. The depth of FZ in the modeling were counted from the top surface to the depth at where T = 1085 K, the melting temperature. The thickness of HAZ was counted from the bottom of FZ to the location at where final crystalline VF goes below 2% (see the inset of Fig. 6). The experimental data were measured from the cross-section images presented before.It can be seen from Fig.

6). The experimental data were measured from the cross-section images presented before.It can be seen from Fig. 10 that for relatively low scan speeds (10 mm/s and 20 mm/s), the HAZ is thicker than 100 μm. At the scan speed of 50 mm/s, the thickness of HAZ is calculated around 50 μm. Further increases of laser scan speed lowers the HAZ thickness to almost 0 at 90 or 100 mm/s of scan speeds. Similarly, the depth of FZ also decreases as the laser scan speed increases, as expected. The depth of FZ at P = 100 W and v = 100 mm/s is 105 μm, which is close to the experimental result, 110 μm, shown in Fig. 4(c). Since at this speed, no crystallized HAZ is observed and, hence, further increase of scan speed can only decrease the depth of FZ.ConclusionsIn this paper, evolution of the crystalline phase in HAZ during laser welding on Zr- based MG is studied.

Using a model couples with heat transfer and classic nucleation theory, the crystallization during laser-MG interaction is accurately predicted. Based on the aforementioned discussion, we can draw the following conclusions: (1) In laser-MG interaction, the major challenge is to suppress the crystallization in HAZ. Increasing laser scan speed will reduce the size of crystallized HAZ. In the HAZ for low laser scan speed (10 mm/s), three distinct crystallized regions was observed. This is explained by the different temperature history induced different crystallization history. (2) Classic nucleation theory is successfully applied in describing the crystallization process of LM105 alloy. The modeled results marches well with the experiments conducted.

(2) Classic nucleation theory is successfully applied in describing the crystallization process of LM105 alloy. The modeled results marches well with the experiments conducted. This model can be used in future work to further studying the laser welding and laser- based additive manufacturing of metallic glasses.767AcknowledgementsThe authors would like to acknowledge the financial support by the Intelligent Systems Center at Missouri S&T, the Department of Energy (Grant Number: DE-FE0012272) and the UM Fast Track Funding Program.References[1] W.L. Johnson, Bulk metallic glasses − a new engineering material, Current Opinion in Solid State and Materials Science 1 (3) (1996) 383-386. DOI http://dx.doi.org/10.1016/S1359- 0286(96)80029-5.[2] E.S. Park, D.H. Kim, Design of bulk metallic glasses with high glass forming ability and enhancement of plasticity in metallic glass matrix composites: A review, Metals and Materials International 11 (1) (2005) 19-27.

nstruction of Open-Source Laboratory-Scale Binder Jetting System for High-Speed Synchrotron X-Ray ImagingJacob Lawrence, Hector Andres Peña Vega, Bryant Stegman, Caleb Roberts, JosephSpencer, Clinton James, McKay Christensen, Nathan CraneDepartment of Mechanical Engineering, Brigham Young University, Provo, UT 84602AbstractAlthough commercial binder jetting (BJ) printers are available, they typically do not allow sufficient control over process parameters needed to study fundamental process characteristics. This work presents an overview of the design and construction of a custom BJ system used to observe fundamental phenomena in the BJ process. CAD models for the design and information on the software of this system is also given. This system will help elucidate the mechanisms that introduce part defects and other challenges unique to the BJ process.

This system will help elucidate the mechanisms that introduce part defects and other challenges unique to the BJ process. The BJ system was designed for both laboratory-scale experiments with a 100 x 100 mm build box and high-speed synchrotron X-ray imaging with a 500 μm thick powder bed, requiring high-accuracy motion stages and a controller with precise timing. The printer includes functionality for depositing and rolling powder, printing multi-layer parts, and direct observation of the jetting nozzle. This BJ system has enabled experiments that provide insight into the printing process that will aid future efforts to mitigate challenges associated with BJ.IntroductionBinder Jetting (BJ) is an additive manufacturing (AM) process that has received added focus in recent years [1]. It utilizes an aqueous binder which is ink-jetted, layer by layer, onto a prepared powder bed to form geometry.

It utilizes an aqueous binder which is ink-jetted, layer by layer, onto a prepared powder bed to form geometry. This process involves powder deposition onto a powder bed, rolling powder to create a smooth surface, ink-jetting binder droplets in precise locations, and lowering the build space for the next layer. These steps repeated create a part that can then be removed from the loose powder and sintered to a final part [1, 2]. Control over each aspect of the operation gives opportunity for precise experiments to better understand BJ.BJ along with other AM processes can manufacture parts with extreme complexity, can vary part shape without additional tooling, and builds parts directly from 3D models [3, 4]. BJ has distinct advantages relative to other AM methods including the separation of the formation step from the sintering process, allowing for greater microstructural control of the final piece [1].

BJ has distinct advantages relative to other AM methods including the separation of the formation step from the sintering process, allowing for greater microstructural control of the final piece [1]. BJ volume production can also be scaled with relative ease for either low or high-volume production of parts. This is accomplished by increasing the number of nozzles used and/or stacking many parts in one build volume. Additionally, BJ can print with nearly any powder feedstock. Testing has been done with copper, stainless steel, sand, and other ceramics [5-8], although metal powder is commonly used. These unique beneficial characteristics make BJ a highly attractive manufacturing process [1]. Although these capabilities are promising, the BJ process requires extensive study due to porosity defects that are introduced during the process as shown by previous research [6, 9].

Although these capabilities are promising, the BJ process requires extensive study due to porosity defects that are introduced during the process as shown by previous research [6, 9]. Research shows that porosity defects in BJ parts form during the printing process by mechanisms such as powder bed flow and spreading, powder particle size, inkjet droplet impact onto the powder bed, and agglomerate formation of binder and powder [10-13]. With further study1506of BJ process, these mechanisms can be more fully understood and printing parameters that minimize defect formation can be determined.Commercial binder jetting machines such as the X1-Lab made by ExOne have been used to study the binder jetting process using high-speed X-ray imaging [14]. These and other commercial systems such as ExOne’s R2 and Innovent systems have been used for other studies [5, 6, 13].

These and other commercial systems such as ExOne’s R2 and Innovent systems have been used for other studies [5, 6, 13]. However, a custom system provides distinct advantages over existing commercial solutions including greater control over printer configuration, capabilities, and user-programmable printing parameters. Open-source binder jetting machines such as the Oasis 3DP and Plan B systems provide low-cost alternatives to commercial solutions but lack high-accuracy motion systems, fine control over printing parameters, hopper-fed powder deposition, and direct observation of printheads for measurement and maintenance.Custom laboratory-scale BJ systems have been constructed for studying the BJ process [15, 16].

The system designed and constructed by Oropeza and Hart [16] includes micron accuracy motion, a heat lamp for curing deposited binder, a single piezoelectric nozzle for control of jetting parameters, a droplet observation system, and a powder bed equipped with a load cell for powder flow measurement and quantification. These capabilities enable numerous controlled experiments to better understand the BJ process [16]. However, the design has not been optimized for outside imaging.

These capabilities enable numerous controlled experiments to better understand the BJ process [16]. However, the design has not been optimized for outside imaging. A printer with an exposed build box during printing is ideal for tests involving an outside source video capture of the process such as high-speed X-ray synchrotron imaging.Outlined in this work is a design for a custom laboratory-scale BJ system [17] with optimization for X-ray imaging, ability to easily configure and program the printer for desired testing, micron accuracy motion, a single-nozzle printhead with control of jetting parameters, hopper-fed powder deposition, and a droplet observation system to quantify droplet features using a stroboscopic effect to capture droplet motion [18].DesignFigure 1.

A CAD model [17] of the custom BJ printer with key components labeled.1507System OverviewThe printer is capable of automated powder deposition, powder spreading, and printing enabling multi-layer experiments to study the BJ process. As seen in Table 1, the printer uses a 100 mm square build plate with 15 mm of travel. The build box can be removed for part curing and extraction. The printer uses two motion stages to position the single nozzle printhead and a third motion stage to move the build platform for multi-layer printing. The printer is equipped with an integrated USB camera for observing and measuring jetting performance of the printhead. The printer is controlled with custom desktop software that runs on a host computer. More information on printer components is available in the following sections. A repository of CAD files for the printer can be found here [17].Table 1.

More information on printer components is available in the following sections. A repository of CAD files for the printer can be found here [17].Table 1. Key printer specifications of the custom BJ system.ParameterValueBuild Volume100 x 100 x 15 mm (L x W x H)Machine Dimensions1040 x 600 x 474 mm (L x W x H)Motor Step Size1.0 x 1.25 x 0.5 µm (X, Y, Z)Maximum Traverse Velocity3000 x 300 x 5 mm/s (X, Y, Z)PrintheadMicroFab Single-Nozzle Drop-on-DemandFigure 2. The custom BJ system configured with the full-sized build box for laboratory testing.1508Linear Motion SystemFigure 3. CAD model [17] of the linear motion system of the printer. It allows movement in all three axes X, Y, and Z.The printer uses three motion stages for positioning the jetting nozzle, build box, and build platform. The X-axis positions the printhead and is mounted above the Y-axis on a bridge. The Y- axis moves the build box assembly and is mounted directly to the baseplate.

The X-axis positions the printhead and is mounted above the Y-axis on a bridge. The Y- axis moves the build box assembly and is mounted directly to the baseplate. The Z-axis actuates the build plate vertically and is mounted to the Y-axis.X-AxisThe X-axis (Parker mSR100) is responsible for positioning the inkjet printhead in the primary jetting direction. The motion stage has 150 mm of travel and a maximum speed of 3000 mm/s. From manufacturer testing, the stage used has a measured accuracy of 3.3 µm and a bidirectional repeatability of 1.6 µm. The X-axis is also used to position the printhead to the rightmost limit of the stage for printhead maintenance and jetting observation.Y-AxisThe Y-axis (Parker 404XR) is responsible for positioning the build box assembly for powder spreading, rolling, and printing. At 500mm in length, the Y-axis can move the build box along most of the length of the printer and is used for powder deposition, rolling, and printing.

At 500mm in length, the Y-axis can move the build box along most of the length of the printer and is used for powder deposition, rolling, and printing. The manufacturer lists a positional accuracy of 34 µm and a bidirectional repeatability of 3.0 µm. The stage uses a Galil BLM-N23 servo motor with rotary encoder that results in a 1.25 µm encoder resolution along the length of the stage.1509Z-AxisThe Z-axis (Newmark VS-15-2) is responsible for positioning the build plate and is used for moving the powder bed during multi-layer printing. The stage is mounted on the Y-axis and attaches to the build plate using a manual actuating gripper to allow for build box removal. More information is available in the section on the removable build box.

More information is available in the section on the removable build box. The stage has 15 mm of travel, and the manufacturer lists an accuracy of 10 µm and a repeatability of 1 µm.Motion ControllerThe three motion stages are controlled by a Galil DMC-4080 motion controller with an AMP-D3547 amplifier module allowing for control of both brushless servos and stepper motors. The motion controller receives commands over ethernet from a host PC which are then processed to move the stages as desired. More information on the custom software used by the host PC to send commands to the motion controller is available in the control software section.Inkjet PrintheadFigure 4. Inkjet printhead assembly. Binder used for printing is stored in a syringe which is mounted to the X-axis above the printhead.

Inkjet printhead assembly. Binder used for printing is stored in a syringe which is mounted to the X-axis above the printhead. The jetting nozzle is mounted in an attachment on a micrometer stage to allow precise changes of the nozzle height along the Z-axis.The printer uses a single nozzle drop-on-demand piezoelectric printhead (MicroFab MJ- AB) controlled by a JetDrive III controller. The printhead is mounted to a manual micrometer stage connected to the X-axis, allowing for positioning the nozzle at a desired height above the powder bed. Binder is stored in a syringe mounted to the X-axis. Backpressure in the binder supply is controlled with a MicroFab pneumatics console.1510Droplet ObservationThe droplet observation system (Figure 5) enables visual confirmation of stable jetting and estimation of droplet velocity. Observation is accomplished using a USB camera (IDS UI-3370CP- M-GL), a zoom lens (Edmund Optics VZM 450 Zoom Imaging Lens), and a strobing LED.

Observation is accomplished using a USB camera (IDS UI-3370CP- M-GL), a zoom lens (Edmund Optics VZM 450 Zoom Imaging Lens), and a strobing LED. The LED is connected to the JetDrive controller and synchronized with the jetting frequency to create a stroboscopic effect to enable visualization of droplets at much lower framerates than traditionally required.Estimating Droplet VelocityDroplet velocity is an important parameter when studying the BJ process. By gradually increasing temporal strobe offsets compared to jetting pulses, the droplet observation camera captures a video depicting a droplet as it exits the printhead nozzle (Figure 6). This video can be used to estimate the droplet velocity by comparing relative droplet positions between frames and calculating how fast it moves between each position. Droplet tracking and velocity estimation is done by a MATLAB script that analyzes the video.

Droplet tracking and velocity estimation is done by a MATLAB script that analyzes the video. Jetting parameters can then be changed, and droplet velocity estimated again until the desired jetting velocity is reached.Estimating Droplet VolumeDroplet volume is another important printing parameter in the BJ process. A volume measurement is conducted by first measuring the mass of a small plastic container. The printer then jets binder droplets at 1000 Hz for three minutes into the container. The container with binder is weighed again to obtain the total mass of droplets jetted during the three minutes. The binder volume is calculated from the measured binder density (1.05 g/cm3). Then the total volume is divided by the number of droplets jetted during the three minutes to obtain an average volume of a single droplet. Average droplet diameter can also be estimated assuming a spherical droplet shape.Figure 5. Droplet observation setup.

Average droplet diameter can also be estimated assuming a spherical droplet shape.Figure 5. Droplet observation setup. The jetting nozzle is positioned between camera and a strobing LED for imaging and maintenance.1511Figure 6. Sequence of binder droplets captured by the droplet observation camera. The velocity of the droplet can be calculated by tracking the droplet position between frames.Powder Deposition and SpreadingThe printer includes a powder hopper and roller for automated powder rolling and spreading. The hopper and the 12mm roller are components used in the ExOne Innovent+. The hopper is powered using a Telsonic SG4L Pro Screening Generator which is controlled via RS- 232 from the Galil motion controller.Figure 7. Cross section view of powder and deposition system. The build box assembly traverses along the Y-axis to position the build box under the hopper and roller for powder deposition and spreading.

Cross section view of powder and deposition system. The build box assembly traverses along the Y-axis to position the build box under the hopper and roller for powder deposition and spreading. When rolling a new layer, the build box first moves to the far right in the image and then travels left under the hopper and roller to receive powder and roll the new layer.1512Removable Build BoxThe printer features a removable build box with a reservoir for excess powder surrounding the build platform. The build box can be removed for part curing and extraction, as well as for removing excess powder after prints for recycling. The build box is attached to the Y-axis using four bolts and the build platform is connected to Z-axis using a manually actuated gripper mechanism. The gripper was designed to ensure that the build platform is aligned properly with both the Z-axis and build box to avoid binding during motion.

The gripper was designed to ensure that the build platform is aligned properly with both the Z-axis and build box to avoid binding during motion. The gripper can be actuated to connect or disconnect the build platform using two fine adjustment screws visible through openings on both sides of the build box assembly.Figure 8. Cross section view of the build box assembly. The build plate is made of a top and bottom plate to accommodate a felt seal that prevents powder from escaping between the build plate and build box. The build plate is mounted to the Z-axis using a manually actuated gripper.High-Speed X-Ray Imaging CompatibilityThe compact design and open format of the custom BJ system enables integration with other equipment for specialized testing. The build box assembly can be removed and replaced for testing that requires a different build box size or configuration.

The build box assembly can be removed and replaced for testing that requires a different build box size or configuration. A custom sample holder mount that attaches to the Y-axis was designed to enable observation of the printing process with high- speed synchrotron X-rays at the Advanced Photon Source in the Argonne National Laboratory. The imaging setup used at the Advanced Photon Source is described in detail by Parab, et al [14].The open design of the printer enabled the X-ray source to pass through the sample during printing without obstruction from other components, and the custom sample holder mount allowed printing in a powder bed thin enough for X-ray transmission. The printer is equipped with a digital output connecter that can be used to trigger external measurement equipment at a user-specified time before the printhead reaches a desired position during printing.

This output trigger was used to synchronize printer operation, high-speed X-ray capture, and optical high-speed image capture during testing at the Advanced Photon Source [19].1513Control SoftwareThe printer is controlled by custom desktop software that handles communications with the Galil motion controller, JetDrive controller, and droplet observation USB camera. All software is available at [17]. The software includes functionality for manual operation of printer motion stages and modules. A user interface is provided for printing primitives which allows the user to set key printing parameters for desired testing. Once testing parameters have been selected by the user and the user selects to start the print, the software sends commands over ethernet to the Galil motion controller which the controller then processes into signals sent to the motion stages to achieve desired movement.

Direct communication with the JetDrive controller over RS-232 is integrated into the software and allows for jetting parameters to be automatically changed during or between prints. The software incorporates an SDK provided by the manufacturer for the USB camera to enable integrated observation of camera footage and control over imaging parameters used for droplet observation. The user interface for observing droplets and setting imaging and jetting parameters is shown in Figure 9.The desktop software was developed in C++ using the Qt5 framework. Communication with the Galil motion controller is handled by the Galil gclib API. Communication with the USB camera is handled by the IDS ueye SDK [17].Figure 9. Screenshot of the custom PC host software used to control the BJ printer.

Communication with the USB camera is handled by the IDS ueye SDK [17].Figure 9. Screenshot of the custom PC host software used to control the BJ printer. The user interface shown is used for observing droplet formation using the droplet observation camera and setting jetting parameters for use during printing.1514System ValidationPowder Deposition and SpreadingPowder deposition and spreading performance of the custom BJ system was measured by preparing a rolled powder bed of stainless steel 316L powder and measuring the packing fraction of the powder bed. The stainless steel 316L powder used had a volume average particle diameter (Dv50) of 10.0 µm (Dv10 of 3.92 µm and Dv90 of 22.0 µm)[15]. The powder bed spread density was measured by rolling a powder bed with 100 layers and using a sharp cylindrical punch with a known inner diameter to isolate a known volume of powder [6].

The powder bed spread density was measured by rolling a powder bed with 100 layers and using a sharp cylindrical punch with a known inner diameter to isolate a known volume of powder [6]. Powder external to the punch was brushed away and the powder inside the punch was weighed to calculate the powder bed spread density. The packing fraction of the powder bed prepared on the custom BJ system was calculated as 55%, indicating that the powder deposition and spreading capabilities of the printer are sufficient to produce parts that can be effectively densified during sintering [1].Printhead AccuracyTo evaluate the performance of the inkjet printing capabilities of the custom BJ printer, a grid of dyed binder droplets was printed on photopaper, and individual measured droplet positions were compared to the commanded positions. The grid was printed with a droplet spacing of 250 µm between each droplet and 20 droplets were printed in each direction to form a 5 mm square grid of droplets.

The grid was printed with a droplet spacing of 250 µm between each droplet and 20 droplets were printed in each direction to form a 5 mm square grid of droplets. Figure 10 shows an image taken on a Keyence VHX-7000 microscope of a printed grid. During testing, the printer starts in the bottom left of the image and scans across each row from left to right to print the grid. The printhead returns to the left side of the grid before each row is printed so each row is printed left to right. The microscope image shown in Figure 10 was processed in ImageJ (FIJI) to calculate the position of the center of each droplet. The grid of measured droplet positions was then aligned to the grid of commanded droplet positions to measure the placement error of individual droplets.

The grid of measured droplet positions was then aligned to the grid of commanded droplet positions to measure the placement error of individual droplets. Absolute error in both the X and Y directions was calculated for each droplet by calculating the difference in position between the droplet in the regular commanded grid and the measured position of the droplet after printing. Error in the X and Y directions was combined to produce Figure 11 which depicts each measured droplet superimposed over the commanded droplet position where measured droplets are color-coded by the total absolute positional error of the droplet.Figure 10. Microscope image of a printed grid of droplets on photopaper. Images were taken on the Keyence VHX-7000 Microscope. The print was conducted with a droplet spacing of 250 µm.1515Figure 11. Color-map of droplet positional error in a printed grid.

Images were taken on the Keyence VHX-7000 Microscope. The print was conducted with a droplet spacing of 250 µm.1515Figure 11. Color-map of droplet positional error in a printed grid. Colored circles show measured positions of printed droplets which are graphed over black outlines of the commanded droplet positions.When aligning the printed grid with the grid of commanded droplet positions for analysis, an iterative closet point (ICP) least squares regression optimization method was implemented. This alignment minimized the total error between all droplets and was used to produce the result shown in Figure 11. The average and standard deviation in absolute error for this test of the X (nozzle movement direction) and Y-axis is 16.3 ± 12 µm and 11.9 ± 9.5 µm respectively.ConclusionsThis custom laboratory-scale BJ system was constructed to better understand the BJ process.

Capable of powder deposition, powder rolling, compatibility with many powder types and sizes, and control over layer thickness allow the powder bed to be customized for studying the BJ process. The high-accuracy motion system enables consistent printing performance for demanding printing conditions. The modular approach of the printer design allows for capabilities to be added for tests that require additional components. Finally, a droplet observation camera provides the capability to observe and measure jetting parameters which enables insight into effects of droplet velocity and volume on the BJ process. Additional features to be added in the future for greater functionality include a heat lamp and a second roller.

Additional features to be added in the future for greater functionality include a heat lamp and a second roller. The current printer and future modules will enable a growing range of experiments the custom BJ system will perform to better understand the BJ process.1516AcknowledgmentsThis material is based upon work supported by the National Science Foundation through award number CMMI-1946724 and NRC 31310019M0006.Disclaimer: Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the views of the National Science Foundation.References[1] M. Ziaee, N.B. Crane, Binder jetting: A review of process, materials, and methods, Additive Manufacturing, 28 (2019) 781-801.[2] M. Li, W. Du, A. Elwany, Z. Pei, C. Ma, Metal Binder Jetting Additive Manufacturing: A Literature Review, Journal of Manufacturing Science and Engineering, 142 (2020) 1-45.[3] I.a. Gibson, D.a. Rosen, B.a.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCONTINUOUS FIBER ANGLE TOPOLOGY OPTIMIZATION FOR POLYMER FUSED FILAMENT FABRICATIONR. Hoglund and D.E. Smith Department of Mechanical Engineering. Baylor University, Waco, TX 76798AbstractMechanical properties of parts produced with the Fused Filament Fabrication (FFF) process are known to be dependent on the printed bead direction, especially when short carbon fiber reinforcement is added to the filament. Given that many FFF filament suppliers now offer carbon-fiber filled products, a unique opportunity emerges in the design of polymer composite FFF parts since bead and fiber direction can potentially be prescribed to give the best structural performance.

As FFF moves from a technology for rapid prototyping and the hobbyist to a viable additive manufacturing method, it is important to also have a design tool that takes advantage of the opportunities that present themselves when polymer composites are employed. This paper presents a topology optimization method for continuous fiber angle optimization approach (CFAO), which computes optimal material distribution (as in the well known SIMP method) in addition to a preferred fiber angle direction by minimizing compliance of statically loaded structures. Our computed results show the effects of variable orientation angle on fiber reinforced microstructure for the topology of two-dimensional FFF parts. Optimal fiber orientations are shown to align with the axis of structural members that form within the structure as expected.

Optimal fiber orientations are shown to align with the axis of structural members that form within the structure as expected. Example design problems are solved and then printed on desktop 3D FFF printers using the material distribution results and a simple infill method which approximates the optimal fiber angle results by a contour-parallel deposition strategy. Mechanical stiffness testing of the printed parts show improved results as compared to structures designed without accounting for the direction of the composite structure. Future work includes extension of the method to three dimensional structures for further application.IntroductionThe use of Fused Filament Fabrication (FFF) has expanded greatly over the past two decades due in part to its low cost and selection of materials. Unfortunately, the reduced mechanical properties of a typical FFF layered structure must be addressed for this technology to reach its full potential in producing engineering products.

Unfortunately, the reduced mechanical properties of a typical FFF layered structure must be addressed for this technology to reach its full potential in producing engineering products. The method of extrusion during part processing has been shown to give the final FFF product non-isotropic properties with increased stiffness in the deposition direction [1]. The stiffness and strength of ABS polymer structures produced with FFF has also been shown to improve with the added reinforcement of single- walled carbon nanotubes and short glass fibers [2], [3].

The stiffness and strength of ABS polymer structures produced with FFF has also been shown to improve with the added reinforcement of single- walled carbon nanotubes and short glass fibers [2], [3]. Recently, the effects of carbon fiber content and fiber length have been found to increase the Young’s modulus of FFF parts to varying degrees [4], increasing the modulus of elasticity by as much as 4X in the print direction.Topology optimization has been shown to be an effective tool for the design of structures produced with additive manufacturing where an optimal material distribution in a design domain is determined for a given material volume [5], [6]. Topology optimization is commonly performed to maximize the stiffness of a part while retaining a given fraction of material volume1 1078in the design domain. The Homogenization Method was developed to determine an optimal material distribution and orthotropic material orientation distribution within a structure [7].

The Homogenization Method was developed to determine an optimal material distribution and orthotropic material orientation distribution within a structure [7]. Unfortunately, the homogenization method, as well as SIMP methods that have been extended to include pointwise material orientation as part of the design problem [8], is impractical to implement in FFF, as it is not possible to produce a part where each point within the structure has a unique material direction and/or continuously varying density. An alternative approach, the SIMP (Solid Isotropic Material with Penalization) method [9], or density method [10], computes material distributions for isotropic materials. The SIMP method has seen extensive use in additive manufacturing where unique three dimensional microstructures designed for a prescribed mechanical response are printed with additive manufacturing methods (see e.g., Hollister, et al. [11]).

[11]). More recently, the distribution of varying microstructure has been computed with a modified SIMP method where the printed optimal design includes the details of optimal unit cell geometries (see e.g., Sundararajan (2010) and Zhang, et al. [12]), [13]. This later research computes the effective orthotropic material matrix for a unit cell having varying density which serves as the design variable in the topology optimization. Alamo and da Silva extended the SIMP method to compute a material distribution for a fixed orthotropic material distribution in a biomechanics structure [14].Unfortunately, none of these methods have addressed the need for having discrete material distributions (such as those computed with SIMP methods (see e.g., [15,16,18]) and also a preferred fiber direction that are suitable for parts produced with the FFF process. This paper builds on our earlier work of topology optimization with a fixed fiber angle [20,21].

This paper builds on our earlier work of topology optimization with a fixed fiber angle [20,21]. Here we consider a modified SIMP approach which we call the Continuous Fiber Angle Optimization (CFAO) method that determines the optimal distribution of an orthotropic material and its orthotropic material orientation to design fiber-reinforced FFF parts. This paper also includes a methodology for printing FFF parts that have the optimal topology and fiber angle, and provides experimental results to verify the usefulness of our CFAO approach.Two-Dimensional Continuous Fiber Angle OptimizationA Continuous Fiber Angle Optimization (CFAO) method is introduced here that is used to compute the optimal topology and fiber angle of planar FFF structures that are produced with an orthotropic microstructure.

The equilibrium equations for the structure are solved with the finite element method as KU = F where K, U, and F are the global stiffness matrix, displacement vector, and applied load vector, respectively. In the above, the compliance is defined as the product U and F, and may be computed at the element level in terms of the element density (cid:2025)(cid:3032), displacement vector ue and stiffness matrix ke as shown.Note that each ke is multiplied by its element density (cid:2025)(cid:3032) which is raised to the power p which is identified as the penalization parameter [9]. The penalization parameter p is fixed with respect to design and causes each (cid:2025)(cid:3032) to move toward either its lower or upper bound, i.e., (cid:2025)(cid:3040)(cid:3036)(cid:3041) or 1, respectively, in the optimization. The element fiber angle array (cid:2242) (with each element fiber angle (cid:2016)(cid:3032) being constrained between (cid:3399)2(cid:2024) in Equ.

Our results have shown that this filter alone eliminates checkboarding without the need for also including a filter on either the (cid:2025)(cid:3032) or (cid:2016)(cid:3032).Our CFAO method is implemented using a custom 2D plane stress finite element program and the fmincon function of Matlab (The Mathworks, Inc., Natick, MA). Isoparametric finite elements are used and the program also computes the design sensitivities from Equs. 8-10 once the finite element solution U is known. The user defined gradient option of fmincon is employed for the compliance objective function, and the volume constraint is implement using the linear inequality constraint feature. The interior-point optimization algorithm in fmincon was used since it accepts gradients and was found to work well with our large scale optimizations.

The interior-point optimization algorithm in fmincon was used since it accepts gradients and was found to work well with our large scale optimizations. In the following examples, f = 0.5, and p = 3, and the initial design was specified as (cid:2025)(cid:3032) (cid:3404) 0 for all density variables and (cid:2016)(cid:3032) (cid:3404) 0 for all fiber angle design variables. The accuracy of the design sensitivities computed with Equs. 8 and 9 was verified with finite difference derivative computations using the forward finite difference method prior to running the topology optimizations that appear below.4 1081(8)(9)(10)CFAO Computational ResultsIn order to illustrate the applicability of the orthotropic model to polymer composite FFF part design, several topology optimization of a 2D model are considered based on the Messerschimitt-Bölkow-Blohm (MBB) beam [17]. The MBB beam is a truss-like structure popularized in many topology optimization studies.

The MBB beam is a truss-like structure popularized in many topology optimization studies. Here we first consider the topology optimization of a half-symmetry model of the MBB beam appearing in Figure 1, and then solve the optimization problem for the complete domain in Figure 3. In all cases, the lesser Young’s modulus is set to 1 while the modulus ratio a ranges from 1 to 10. Poisson’s ratio of the major direction (greater Young’s modulus) is defined to be 0.36, an approximate ratio for polylactic acid (PLA) polymer-printed structures, and p = 3. The topology optimization scheme iterates until the greatest change in elemental density is less than 1%. In the following, all elements contain a density and fiber angle design variable which produces a topology optimization problem from Equs. 1 and 2 where the number of design variables equates to two times the number of elements.

1 and 2 where the number of design variables equates to two times the number of elements. The thickness of the model into the plane is h = 1.We first consider the simple cantilever beam design shown in Fig. 1 to illustrate our CFAO topology optimization process. The beam finite element model has 30 elements in the horizontal direction and 10 elements in the vertical direction. The elastic constants are E1 = 5 and E2 = 1, and 12 = 0.3, making  = 5. The applied load is 1 and acts downward on the right lower edge of the design domain.Figure 1: Cantilever beam example showing loading and boundary conditions.Selected designs taken from iterations of the CFAO topology optimization appear in Fig. 2. Elements in the figure appear dark as (cid:2025)(cid:3032) approaches unity (i.e., element full of material), and white as (cid:2025)(cid:3032) approaches zero (i.e., no material). Also shown in each element is an arrow indicating the fiber angle direction.

Also shown in each element is an arrow indicating the fiber angle direction. Note that all (cid:2016)(cid:3032) are zero for the initial design and quickly change to become parallel to the direction of the dominant structure in the truss-like members. The topology of the structure emerges after just 25 iterations, however, the edge of the truss-like becomes more distinct as the optimization concludes. We found that the minimum compliance indicated a much stiffer structure when compared to our earlier results that only considered a variable element density while fixing the fiber angle in either the 1- or 2-direction.We next consider the MBB beam design space appearing in Fig. 3. The design domain has a base width of 120 and a height of 20, yielding 2400 elements and 4800 design variables in the CFAO optimizations. Material properties and optimization initial conditions are the same as those given above for the cantilever beam.

Material properties and optimization initial conditions are the same as those given above for the cantilever beam. The load is 1 unit in the negative x2 direction applied5 1082at the midpoint of the top of the beam. Optimal topologies for varies values of (cid:2009) ranging from 1.33 to 10 appear in Fig. 4. Notice that little change in topology is seen for (cid:2009) > 1.33 indicating that the modulus ratio has little effect on topology when fiber angle is included as a design variable in the optimization.Figure 2: CFAO results for a cantilever beam after a) 0 iterations (i.e., initial design), b) 5 iterations, c) 25 iterations, and d) final iteration 48.Figure 3: MBB beam design space with boundary conditions and applied load.6 1083Figure 4: CFAO optimization results for the MBB beam with varying (cid:2009).

a) (cid:2009) = 1.33, b) (cid:2009)= 1.5, c) (cid:2009) = 2, d) (cid:2009) = 5, and e) (cid:2009) = 10.Values of optimal compliance, CPU time, and number of optimization iteration are given in Table 1. It is shown that the optimal compliance decreases considerable, however, this value alone is not a good comparative measure between these designs since the overall stiffness of the material increases with (cid:2009) since E2 = 1 for all optimizations. It is interesting that as (cid:2009) increases, the number of optimization iterations do so also, as does the CPU time (CPU time reported for an Intel i7-4930K CPU at 3.4 GHz and 64 GB RAM). CPU times for the CFAO optimizations7 1084where found to be significantly higher than our earlier fixed fiber angle optimizations as described in Hoglund [21].

CPU times for the CFAO optimizations7 1084where found to be significantly higher than our earlier fixed fiber angle optimizations as described in Hoglund [21]. This is due, in part, to the use of different optimization algorithms for the CFAO and fixed angle optimizations which employed the optimality criteria method given in Sigmund [15].Upon comparing these results to our earlier work with fixed horizontal fiber angle (i.e., all (cid:2016)(cid:3032) (cid:3404) 0) [20,21], we find that a lower minimum compliance is obtained for all (cid:2009) (cid:3408) 2 using the CFAO method. Optimization problems are inherently nonlinear, which can result in convergence to various local minimum points. Our results tended to converge to topologies typical of earlier MBB beam optimization, mitigating our concern that local minimum is an issue.

Our results tended to converge to topologies typical of earlier MBB beam optimization, mitigating our concern that local minimum is an issue. Future work should include additional computations to ensure that solutions obtained here are indeed the global minimum.Modulus ratio (cid:2009)1.33 1.5 2 5 10Number of Iterations 44 62 72 180 169CPU Time (sec)586 911 994 2379 2224Minimum Compliance 86.21 70.39 53.59 22.54 12.06Table 1: CFAO MBB optimization results for various modulus ratio (cid:2009).Fabrication and Testing of Optimal CFAO DesignsIn order to further demonstrate the usefulness of the CFAO method, optimized MBB beams are printed with varying orientations using carbon fiber-reinforced PLA sold as 3DXMax CFR-PLA by 3DxTECH (3DXTECH, Byron Center, MI). The filament contains approximately 15% carbon fibers by weight in a 1.77 mm diameter filament and has a published tensile strength of 47.9 MPa and a tensile modulus of 4.791 GPa.

The filament contains approximately 15% carbon fibers by weight in a 1.77 mm diameter filament and has a published tensile strength of 47.9 MPa and a tensile modulus of 4.791 GPa. For FFF processing, our MBB structures are printed with a MakerBot Replicator 2.0 3D FFF printer with a hardened steel 0.6 mm diameter nozzle. The geometry of the printed MBB beams was taken from the optimized result for  = 10 that appears in Fig. 4f. Our FFF structures are printed with 100% infill to reduce interbead voids within the material, and are printed with a thickness of 5.5 mm and a layer thickness of 0.2 mm. The CFAO MBB beam selected for printing is shown in Fig.

The CFAO MBB beam selected for printing is shown in Fig. 5a which was computed with 180 elements in the horizontal direction and 30 elements in the vertical direction (resulting in 5400 elements and 10,800 design variables).To produce a printable format of the topology optimizations, the code Top3dSTL_v3.m was used to convert the physical array of elemental densities to .stl file format [19]. We use Top3dSTL_v3.m to translate the optimal density distributions to an STL where each element represents a unit cube. The 2-dimensional information for the planar MBB beams is retained while the thickness out-of-plane is modified to adjust for scaling of the part dimensions in the plane. Autodesk MeshMixer (AutoDesk, San Rafael, CA) is used to smooth the edges of the CFAO structures prior to printing.

Autodesk MeshMixer (AutoDesk, San Rafael, CA) is used to smooth the edges of the CFAO structures prior to printing. This smoothing helps to facilitate a printed part that has beads8 1085with a primary orientation that is parallel to the edge of the truss-like members that result from the CFAO optimizations. To realize print paths that remain mostly parallel to the edge of the structure, a contour-parallel deposition strategy is employed. In this case, the strategy is implemented by specifying a high number of shell layers in the MakerBot Desktop software to ensure so that all printed beads approximately align with the edges, particularly within the structural members that form. Fig. 5 shows the model data (left half of beam only) at various stages during the processing of data for printing which includes the CFAO optimized topology and fiber angle, the smoothed image obtained with Autodesk MeshMixer, and the bead pattern print path as specified by the MakerBot Desktop.

The resulting printed beam appears in Fig. 6.a)b)c)Figure 5: CFAO optimized MBB print data (left half only) of a) optimized density and fiber angle distribution (5400 element model), b) MeshMixer smoothed STL file data, and c) MakerBot Desktop print bead pattern.9 1086Figure 6: Printed CFAO optimized MBB beam.Printed CFAO optimized MBB beams were tested in 3-point bending as shown in Fig. 7 using an Instron (Instron, Norwood, MA) tensile test machine with a 2kN load cell. The load is applied to the top center of the printed beam in the same manner as that used in the CFAO optimization. The structural stiffness was computed from force and displacement data that was collected using the Bluehill 3 (Instron, Norwood, MA) software. Loading was performed in the linear force-displacement range for the sample and testing was ended prior to sample failure.Three samples were printed and tested to obtain the stiffness results given in Table 2.

Also shown in Table 2 are results from similar tests (3 samples each) where optimal topologies were computed assuming a fixed horizontal bead path (i.e., (cid:2016)(cid:3032) (cid:3404) 0) and a fixed vertical bead path (i.e., (cid:2016)(cid:3032) = 90o) as described in [20,21]. Note that in these structures with a fixed fiber angle, each element has a single density design variable, which simplify the calculations and the process for creating the print file. The fixed angle optimizations provide a point of comparison to illustrate the effectiveness of the CFAO method. As shown in Table 2, samples optimized and printed with a horizontal bead are 15.5% stiffer than samples optimized and printed with a predominant vertical bead. This result is to be expected since horizontal stiffening is more efficient than vertical stiffening under the given bending load. Table 2 also shows that CFAO optimized MBB beams perform better than either horizontal or vertical optimized and printed beams.

Table 2 also shows that CFAO optimized MBB beams perform better than either horizontal or vertical optimized and printed beams. It is seen that the CFAO beam is 29.9% stiffer than the vertical optimized and printed structure and 12.4% stiffer than the horizontal optimized and printed structure.Figure 7: Mechanical test set up for CFAO optimized MBB beam.More information on details of the optimization and testing may be found in Hoglund[20,21] which also includes the CFAO optimization and testing of the Mitchell truss [17].10 1087Conclusions and Future WorkThe CFAO method for topology optimization of 2D anisotropic structures has been developed which includes both material distribution and angle of preferred modulus, where special consideration is given to carbon fiber-reinforced Fused Filament Fabrication products.

The stiffness of the structures are maximized for a given volume fraction by determining the optimal element densities and fiber angle with a gradient-based optimization algorithm. Adjoint design sensitivities are used to efficiently compute the design gradients. Various topology optimizations are performed using the MBB beam to illustrate our computational approach. Results show that once the modulus ratio exceeds about two, little difference is seen in the optimal topologies of the beam structure, and that the minimum compliance decreases with increase modulus ratio, as expected. The results also indicate that the optimal fiber angle generally aligns with the axis of truss members that form which has been seen by other researchers. A method is also presented for obtaining print path information from CFAO results. The resulting design are printed and tested and show that printed CFAO structures exhibit a higher stiffness than structures optimized for density alone.

The resulting design are printed and tested and show that printed CFAO structures exhibit a higher stiffness than structures optimized for density alone. In summary, the CFAO method presented here is shown to provide optimal topologies and fiber angle for planar structures produced with carbon fiber-reinforced FFF structures.Future work will extend the method given here to included structures with multiple loading conditions and three-dimensional design domains. The results appearing in this paper that include planar structures having a single load set is shown to produce fiber angles that mostly align with truss-like members that form during the topology optimization. Multiple load sets would likely yield quite different results, and may require that special attention be given to the element fiber angle, and how each element’s angle varies with respect to its neighbor. It will be interesting to see if a fiber angle filtering method is required to obtain a printable solution.

It will be interesting to see if a fiber angle filtering method is required to obtain a printable solution. The extension of this approach to three-dimensions is also of interest where the added complexity will result from having non-planar design domains. Again, the continuity of fiber angle will be important here as well. The successful extension of this methodology to three dimensions is expected to make it possible to significantly impact FFF structural design.MBB SampleAverage (N/m)Standard Deviation (N/m)Coefficient of Variationhorizontal bead [20,21] vertical bead [20,21]2.90x102.51x10550.12x1055 0.052x100.0420.021CFAO3.26*1055 0.015x100.0047Table 2: Measured stiffness (average of 3 samples each) of printed CFAO optimized beams with other horizontal ((cid:2016)(cid:3032) (cid:3404) 0) and vertical ((cid:2016)(cid:3032) = 90o) beams (cf. [20,21]) included for comparison11 1088AcknowledgementsThe authors would like to thank the Dr.

[20,21]) included for comparison11 1088AcknowledgementsThe authors would like to thank the Dr. David Jack and the SIC’EM group at Baylor University for their collaboration and assistance, as well as the Baylor University Society of Plastics Engineers for their support.References[1] A. Bellini and G. Selcuk, “Mechanical characterization of parts fabricated using fuseddeposition modeling,” Rapid Prototyp. J., vol. 9, no. 4, pp. 252–264, 2003. [2] M. L. Shofner, K. Lozano, F. J. Rodríguez-Macías, and E. V. Barrera, “Nanofiber-reinforced polymers prepared by fused deposition modeling,” J. Appl. Polym. Sci., vol. 89, no. 11, pp. 3081–3090, Sep. 2003.[3] W. Zhong, F. Li, Z. Zhang, L. Song, and Z. Li, “Short fiber reinforced composites for fuseddeposition modeling,” Mater. Sci. Eng. A, vol. 301, pp. 125–30, 2001.[4] F. Ning, W. Cong, J. Wei, S. Wang, and M. Z.

ntinuous ink-jet printing electronic components using novel conductive inks Junfeng Mei, Michael Lovell, Marlin Mickle, and Steve Heston University of PittsburghReviewed, accepted September 1, 2004AbstractTo reduce the cost of electronics fabrication and to take advantage of numerous potential specialized applications, a novel process of manufacturing three-dimensional electronic products is introduced in the present investigation. This process, which is currently being patented by the University of Pittsburgh, utilizes specialized continuous ink jet (CIJ) printing technology and innovative conductive inks that are capable of producing high-resolution conductive traces. Unlike drop-on-demand (DOD) techniques that are aimed at more precise features sizes (1 micron or less) and smaller scale applications, the present investigation demonstrates that the new process can be used in mass production applications where larger feature sizes (~50 microns) are sufficient.

It has been found that the traces produced by the process have excellent adherence and have an electrical resistivity of only 2.9 times of bulk silver after curing. The major advantage of the proposed CIJ process is that it not only provides a fast and cost-effective method for applying electronic components (conductors, diodes, capacitors and resistors) on existing products, but it also allows the printing of conductive traces in three-dimensional space. With a throw distance that exceeds 10 cm, it demonstrates how the new fabrication process is not only suitable for desktop microfabrication, but also for large volume applications such as automotive glass.IntroductionAdvances in photolithography techniques have allowed the electronics industry to miniaturize components over the past several decades.

Based on these advances, new technological challenges have emerged because of the limited geometric flexibility, increased processing complexity, increased processing time and greater cost of existing fabrication methods. In today’s global marketplace, therefore, there is a strong need to introduce innovative fabrication techniques that can produce high quality products using processes that are both faster and cheaper than traditional methods. In this light, inkjet printing technology offers several potential advantages over direct write fabrication processes including low capital equipment costs, high materials efficiency, rapid deposition rates, and increased flexibility in the creation of geometric features. In the electronics industry, for example, inkjet printing of circuit interconnects is an additive process where material is only deposited on specified locations, thereby minimizing the amount of material utilized in the circuit board formation process.

In fact, in a fully integrated inkjet system, an entire circuit trace patterns could be designed using computer graphics software and printed directly onto a substrate. This is in contrast to conventional screen-printing techniques that require sophisticated masks. Inkjet printing techniques, therefore, offer a higher degree of automation that will help minimize cost and increase repeatability and reliability.334There are two different ink-jet techniques that have become mature in the printing industry: drop-on-demand (DOD) and continuous ink jet (CIJ). In a DOD system, single drops are ejected through an orifice at a specific point of time. To maintain accuracy, a DOD printing head must be placed vertically above and very close to a substrate (within 1mm). Continuous inkjet systems, in contrast, utilize a pressurized fluid stream and a piezoelectric element to form droplets.

Continuous inkjet systems, in contrast, utilize a pressurized fluid stream and a piezoelectric element to form droplets. By oscillating the piezoelectric element at high frequencies (in the range of 20 to 80 kHz), capillary waves are generated within the fluid domain to break up the fluid stream (fig.1). By carefully adjusting the voltage and frequency of the PZT device, a continuous and consistent stream of fluid droplets of uniform size and spacing can be generated. By placing conductive material in the fluid, an electric charge can be impressed on selected drops as they break-off from the fluid stream at the center of a charge channel as shown in the figure 1. These charged droplets can then be deflected by means of high voltage deflector plates to form various patterns onto the substrate.

These charged droplets can then be deflected by means of high voltage deflector plates to form various patterns onto the substrate. The uncharged droplets are captured by a gutter mechanism and re-circulated through the system.The flexibility, precision, and speed of the CIJ printing process has allowed it to be used in numerous applications within the coding, marking, and labeling markets, as well as the textiles and micro-manufacturing industries [1,2]. CIJ offers benefits over DOD with respect to image quality, throw distance, and the ability to deflect droplets independent of gravity [3,4].

CIJ offers benefits over DOD with respect to image quality, throw distance, and the ability to deflect droplets independent of gravity [3,4]. These advantages make CIJ capable of printing components in three-dimensions (figure 2), which has been the focus of our research at the University of Pittsburgh for the past several years [5].Figure 1 Continuous ink-jet printing: A multiple-deflection system [10]Figure 2 Three-dimensional pigmented CIJ printing in coding and marking (resource: www.linx.co.uk)335Figure 3 PENI tagInkjet printing is most commonly utilized in desktop office printers, but there are numerous other applications that can incorporate ink-jet printing technology. Electronics and biomaterials are the two fields to offer significant potential for utilization of the inkjet printing process. Precision micro dispensing based upon ink jet technology has been used in biomedical applications since the early nineties.

Precision micro dispensing based upon ink jet technology has been used in biomedical applications since the early nineties. It is a key enabling technology in the development of Bio- MEMS devices, bio-molecular sensors, micro-fluidic devices and micro-optical systems [6]. The printing of sensors on medical components, for example, can be tailored to meet specific requirements. The market for manufacturing biomedical sensors has been estimated to exceed several billion dollars [7]. Low-cost RFID tags that are expected to replace UPC barcodes on consumer products, likewise, can be printed using CIJ technology to achieve enhanced consumer convenience and warehousing efficiency through a realization of real-time controls, automated inventory processes and automated checkout [8]. Recent research by the investigators on small antennas has made it possible to fabricate the antenna directly onto the chip as a part of the CMOS fabrication process (fig.3).

Recent research by the investigators on small antennas has made it possible to fabricate the antenna directly onto the chip as a part of the CMOS fabrication process (fig.3). Such a change in the design removes the cost of assembling a separate antenna. Microsystem devices such as resonant inductive coils, electrostatic-drive motors, and actuators can be constructed with ink-jet printing methods [9].Current Conductive InksA major challenge in utilizing inkjet processes for printing electronic components is the formulation of inks with suitable physicochemical properties. Specifically, the rheology properties of the conductive inks must be compatible with available inkjet printing technology. The surface tension and viscosity of the ink determines the ejected droplet velocity, size, stability and the shape of the droplets impinging on the substrate [10].

The surface tension and viscosity of the ink determines the ejected droplet velocity, size, stability and the shape of the droplets impinging on the substrate [10]. The impingement shape establishes trace resolution and thickness, which ultimately determine the trace mechanical and electronic properties. Inks possessing a surface tension on the order of 25-70 mN/m and Newtonian viscosity of 1-10 mPa have shown to be most suitable for use with the CIJ processes (Table 1 [11]). Most commercially available conductive inks possess viscosity and surface tension that exceed levels which allow droplet formation in the range of 30-100 µm size using ink jet technology [12].To satisfy the requirements of inkjet printing, the inks must contain the appropriate precursors and carrier components. Various binders, dispersants, and adhesion promoters are generally included in the composition, depending on the nature of the precursor and the particular application.

Various binders, dispersants, and adhesion promoters are generally included in the composition, depending on the nature of the precursor and the particular application. In the case of inks for metallization, the content of the metallic ink must be adjusted to provide the required resolution, adhesive and electronic properties for the conducting lines. Conductive inks are based on noble metals because of their chemical inertness336and electrical conductivity [13]. There are several categories for ingredients of conductive inks including nanoparticles, microparticles, and metal organic decompositions.Table 1 Typical ink properties for a continuous ink-printer [11]Viscosity (cps) Surface tension (dynes/cm) Particle size1-10 25-70 Micrometer, no agglomerate1. NanoparticlesNanoparticles typically measure 1 to 100 nm in diameter and consist of clusters of metallic, semiconducting, or insulating atoms.

NanoparticlesNanoparticles typically measure 1 to 100 nm in diameter and consist of clusters of metallic, semiconducting, or insulating atoms. Because of their small size, nanoparticles exhibit melting temperatures as low as 1000oC below their bulk material [14]. The lower melting point is a result of their comparatively high surface-area-to-volume ratio, which allows soldering bonds to take place between neighboring particles at relatively low temperatures. Nanoparticles can be sintered at plastic-compatible temperatures as low as 300oC to form a material that is nearly indistinguishable from the bulk material. Their low melting temperature is a major reason that gold and silver nanoparticles can be used as a precursor for conductive ink. Unlike micro scale particles, nano-size particles will not clog a typical 30 µm inkjet orifice, making them capable of producing high-resolution components.

Unlike micro scale particles, nano-size particles will not clog a typical 30 µm inkjet orifice, making them capable of producing high-resolution components. The Media Laboratory at MIT has made conductive inks by dispersing 5-7 nm diameter silver particles that are 10% by weight in alpha-terpineol into a colloid. The printed silver was found to have a resistivity of approximately 3uΩ.cm, which is approximately the twice that of the bulk silver (1.6uΩ.cm) [9]. Researchers at University of California Berkeley have reported a conductive ink of thiol-encapsulated 1.5 nm gold nanoparticles that were dissolved in toluene or alpha-terpineol [15,16].2. Metal Organic Decomposition (MOD)Metallo-organic crystals can be dissolved in an evaporable organic solvent such as toluene, ethanol or butanol to form conductive inks. Metal films can form upon solvent evaporation and decomposition of the metallo-organic compound during curing.

Metal films can form upon solvent evaporation and decomposition of the metallo-organic compound during curing. In the process, gaseous byproducts of the decomposition leave the system, providing contamination-free metal films. Computer controlled ink jet assisted metallization of the grid pattern of solar cells with metallo-organic decomposition (MOD) silver inks offers a maskless alternative method to conventional photo-lithographic thin film technology and screening technology [17]. By curing at a low temperature (350oC) after direct printing, a uniform line film on rough surface solar cells has been found to be a low-cost, fine-resolution, and convenient process. The National Renewable Energy Laboratory reported the inkjet printing of metal organic decomposition (MOD) inks that had good adhesion and measured a contact resistance of 400 µΩ.cm2 [18,19].3.

The National Renewable Energy Laboratory reported the inkjet printing of metal organic decomposition (MOD) inks that had good adhesion and measured a contact resistance of 400 µΩ.cm2 [18,19].3. Micro-scale metal particlesUS patent 6,036,889 disclosed the formulation of a conductive ink formed of a mixture of metal flakes and metallo-organic decomposition (MOD) compounds in an organic liquid vehicle [20]. The mixture contains metal flake with a ratio of the maximum dimension to the minimum dimension of between 5 and 50. The MOD compound begins to decompose at a temperature of337approximately about 300oC to promote consolidation of the metal constituents and bonding to the substrate in a time less than six minutes. The addition of silver flake to the MOD compound allows thicker deposition without loss of resolution.

The addition of silver flake to the MOD compound allows thicker deposition without loss of resolution. By adding silver colloidal with a mean particle diameter of approximately 20 nanometers, the temperature for consolidating the metal has been further reduced to 270oC [20]. Heating the deposited material to temperatures just above about 300oC results in consolidation of the material to near theoretical density with electrical resistivity of around twice that of the bulk metal and good mechanical properties. The current limitation of the micro-scale metal particles is that it is too thick to be inkjet printed with standard techniques.4. Aqueous solution of silver inorganic compoundNano scale particles or MOD precursors are two general resources for synthesizing conductive inks whose rheology properties permit inkjet printing. They work well under small- scale experimental conditions, but resources are expensive and sometimes not available commercially.

They work well under small- scale experimental conditions, but resources are expensive and sometimes not available commercially. In the current research, the investigators have developed a novel aqueous ink composition for CIJ printing that can be broadly applied. The precursors are commercially available and substantially less expensive than metallo-organic compounds. Due to its low cost, silver nitrate was chosen as the primary inorganic compound in an aqueous ink composition. Silver nitrate is used in almost all processes for producing silver compounds and has a wide range of applications that vary from painting, xerography, chemical electroplating, electric batteries and medical catalysts. Silver nitrate melts at 212oC and decomposes to silver at temperatures from 440oC to 500oC. It also has solubility of 219 g in 100gram water, which is significant when very dense ink is needed for high-resolution electric circuits.

It also has solubility of 219 g in 100gram water, which is significant when very dense ink is needed for high-resolution electric circuits. Considering the broad resources and physical properties, silver nitrate makes an excellent candidate if it satisfies other basic prerequisites for conductive inks. Silver chloride is another important silver compound, but it has a high melting temperature of up to 455 °C and is only slightly soluble in water, making it unsuitable for the conductive inkjet printing.Applied Inkjet printing systemThe investigators have constructed a specialized continuous ink-jet system for depositing conductive ink materials. Originally designed and developed at the Swanson Center for Product Innovation (www.engr.pitt.edu/SITE/SCPI), this modular system can produce conductive traces on glass and polymers.

Originally designed and developed at the Swanson Center for Product Innovation (www.engr.pitt.edu/SITE/SCPI), this modular system can produce conductive traces on glass and polymers. Overall, the system consists of three major components: (1) a specialized print head, (2) a control system, and (3) a dual-axis precision positioning table.As shown in Figure 4, the constructed print head includes all of the standard mechanical components required for continuous ink-jet printing – a PZT can, a sapphire orifice, a charge tunnel, deflector plates, and a gutter. In the print head (top view), the charge tunnel is 6 mm in diameter, the deflector plates have dimensions of 30 mm by 10 mm, and the sapphire orifice was 60 µm in diameter. In addition to these mechanical components, the figure shows the underside (bottom view) of the print head, which contains a PCB board and all of the electronic circuitry required to control and generate the droplets.

In the print head, the PZT can was 0.8 inches in diameter and included a 0.2 in-diameter PZT element. The fluid stream being deposited by the print head is controlled using the control system depicted in Figure 5a. The control system338contains a reservoir for the material being deposited and allows direct variation of numerous system variables. These variables include the fluid system pressure, the material flow rate, the voltage applied to the PZT crystal (controls deflection), the vibrating frequency of the PZT crystal, and the electrical properties of the charge tunnel and deflection plates. The final component of the current experimental CIJ system is a dual-axis, precision motion table. The positioning table, which is depicted in the experimental set-up shown in Figure 5b, is computer programmable and has an accuracy of 100 nanometers.Figure 4. CIJ printing head: A - Orifice; B – Charge tunnel; C – Deflection tunnel; D – Gutter(a)(b)Figure 5.

CIJ printing head: A - Orifice; B – Charge tunnel; C – Deflection tunnel; D – Gutter(a)(b)Figure 5. (a) CIJ printing control system; (b) Motion table.Experimental procedure and results of novel aqueous conductive inkPrimary experiments were carried out by printing an aqueous solution of 0.5 gram silver nitrate diluted in 1 ml pure water. After being printed on a glass substrate and cured at 500°C for 15 minutes, the structure shown in Fig 6 was obtained. As demonstrated by the figure, the trace was neither aesthetically appealing nor functional. After analysis, the white thick film in the figure was found to be metallic silver and the black net-shape material was determined to be the metal oxide. The decomposition was observed to be so violent that small drops were exploded and dispersed nearby. When heated above 212°C, the surface tension of the liquid silver nitrate increased significantly and the slim trace broke and contracted to separate spots on the glass substrate.

When heated above 212°C, the surface tension of the liquid silver nitrate increased significantly and the slim trace broke and contracted to separate spots on the glass substrate. The final cured trace was therefore discontinuous and had no real application as an339electric conductor. The adherence of the solid spot to the glass substrate could be easily peeled off, further showing the inability of a simple silver nitrate solution to form a smooth continuous conductive trace. Additives were deemed necessary to change the characteristics for the applications of interest. Specifically, the additives were needed to modify the aqueous silver nitrate to address the following requirements:a. To slow down the decomposition process of the silver nitrate and make itcontrollable;b. To adjust the surface tension of the melted silver nitrate at high temperatures; c. To bond metal silver thick films with glass substrate inherently; d.

To adjust the surface tension of the melted silver nitrate at high temperatures; c. To bond metal silver thick films with glass substrate inherently; d. To be chemically and physically compatible with a silver nitrate water solution.To improve the properties of the silver nitrate solution, numerous chemical additives were selected, tested, and compared based on their physical and thermodynamics properties in the solution together with the silver nitrate. Some additives were found to be effective in improving the conductive silver nitrate ink. In this manner, a viable conductive ink was produced by combining 25~100 gram of silver nitrate, 10~25 gram of additive chemicals and 100 ml of pure water into a transparent aqueous solution. The material was then printed using the CIJ system onto micro glass slides heated to 75oC. The heated glass substrate was found to increase the evaporation rate of the ink, thereby decreasing the width of the trace produced.

The heated glass substrate was found to increase the evaporation rate of the ink, thereby decreasing the width of the trace produced. It is important to note that substrate temperatures above 100oC were avoided because a porous trace structure developed from boiling. The printed solution was dried and crystallized on the glass slide. The glass slides were then put into a stove that was preheated to 485oC. After 20 minutes curing, the silver nitrate solution homogeneously decomposed into a shining metal trace. The color of golden yellow was observed from the opposite side of glass, which indicated intimate chemical bonding of the printed traces onto the glass substrate.Figure 7 shows a SEM structure of the conductive trace that was produced after curing. In the figure, there are many ‘cell-like’ substructures composed of silver conductive film. The silver has a purity of 95% and produces good electric conductance.

In the figure, there are many ‘cell-like’ substructures composed of silver conductive film. The silver has a purity of 95% and produces good electric conductance. There are openings for the metal silver cells, which assure silver conductivity between cells. This property is crucial for continuous conductive traces in a macro-scale because the slightest amount of discontinuity will make the trace nonconductive as a whole. The durable metal oxide ridge surrounds the silver cells to protect the soft silver film from wear and also allows it to strongly adhere to the glass substrate. This special functional composite structure explains traces’ beneficial characteristics. In fact, the pure silver in the developed compounds did not naturally produce a strong adherence with glass.

This special functional composite structure explains traces’ beneficial characteristics. In fact, the pure silver in the developed compounds did not naturally produce a strong adherence with glass. The interlayer of oxide, therefore, should be considered a necessary component for attaining a strong adherence to the glass substrates.theAs listed in table 2, the novel ink has surface tension and viscosity properties that are similar to pure water. The surface tension can be adjusted in the range of 25~70 dynes/cm to satisfy CIJ printing demands and without affecting the final trace characteristics. Compared to the ink properties for the continuous inkjet printers listed in table 1, the new ink is fully compatible with current CIJ technology.

Compared to the ink properties for the continuous inkjet printers listed in table 1, the new ink is fully compatible with current CIJ technology. It is noteworthy to mention that there is no clogging problem during the inkjet printing process after several days of continuous operation.340Figure 6 The SEM photo of decomposition debris after silver nitrate was thermal decomposed on glass slide.Table 2 Fluid properties of normal and dense ink applied experiments.NormalDenseDensity (g/ml) Viscosity (cps) Surface tension (dynes/cm)1.4 1.47 681.68 2.58 82Figure 7 Microstructure of conductive trace after novel conductive ink was cured at 485oC for 20 minutes.ResultsOnce the preliminary ink solution was obtained, a factorial design was carried out to determine the critical processing parameters for optimizing the conductive inks. Based on the factorial design, optimized ingredients of 0.2 gram of additives with 1.0 gram of silver nitrate in 1ml water were determined.

Based on the factorial design, optimized ingredients of 0.2 gram of additives with 1.0 gram of silver nitrate in 1ml water were determined. These ingredients were sufficiently stirred until all the components were completely dissolved and a transparent aqueous solution was achieved. The aqueous solution was then printed with the CIJ system onto a micro slide glass substrate. The glass substrates were heated on a hot plate to 75oC and then cured at 485oC for 20 minutes.341The electric resistance of the optimized ink was measured to be between 0.4~0.6 ohm per inch. As measured by a Surtronic 3+ profilometer, for a typical single-pass printed conductive trace, the trace thickness was measured to be 1.8µm and the width was 1.3mm. The resistance was 0.5 ohm per inch length and the electric resistivity was calculated only 2.9 times of bulk silver ( . The sheet resistance was 0.0256 ohm/square in this case. The silver conductive trace was firmly bonded to the glass substrate.

The sheet resistance was 0.0256 ohm/square in this case. The silver conductive trace was firmly bonded to the glass substrate. Abrasion tests were carried out using a pin-on-disk tribometer with the parameters set as: Abrasion tool: ½ inch (Diameter) steel ball Tip loading: 125g, 250g, 500g Running circle radius: 15 mm RPM: 75 After the pin-on-disk wear experiment, no abrasion of the trace was observed after 1000 cycles under a load of 125g when viewed under a microscope (Fig 8). In-time monitoring of electric resistance along both sides of cutting point during abrasion tests further confirmed the strong adherence of conductive trace to glass substrate (Fig 9.) For a 125 gram tip loading, abrasion was observed to start only after 2500 cycles and the trace was broken after 5000 cycles. The conductive trace withstood up to 1600 cycles when the loading was kept as high as 500g.

The conductive trace withstood up to 1600 cycles when the loading was kept as high as 500g. The results from the abrasion test demonstrated that the current ink can be safely applied in the situations where frequent scratching takes place on the surface of conductors, for example, windshield heaters.

In fact, the abrasion results far exceed current industry standards.8 E −6.4ohmmeter)Figure 8 Abrasion was not observed on the conductor printed with current ink after 1000 runs (Left); Abrasion of the conductor with the ink printed using glass frit as adherence promoter after 300 runs (Right).1.801.60Figure 9 Electric resistance change during 5000 runs of abrasion test, tip load: 125 g, 250g and 500g, respectively.h c n1.40ir e p e c n a t s s e r ci1.201.000.80i r t c e l E0.60tip loading 125g0.40tip loading 250gtip loading 500g0.200.003420100020003000Num of runs for abrasion test40005000ApplicationsOnce the aqueous conductive ink was generated, the continuous ink jet system developed by the investigators was used to produce conductive traces on glass substrates.

The investigators found that uniform size and constant ink droplets could be generated and controlled without satellites disturbance (figure 10) by adjusting the fluid pressure, the voltage and frequency of the PZT crystal, and the charge tunnel pulse. Droplet diameters produced were in the range of 100µm and the droplet velocity was approximately 4m/s. This material was found to possess good viscosity and surface tension properties and was able to produce conductive traces (Figure 11) similar to windshield heating grids after heating at 485oC for 20 minutes. The resolution of the grid lines were approximately 0.4 mm and the printed sample produced 0.5 watts per square inch at approximately 3.5 volts, which was well within the required range for a windshield.As a preliminary experiment to determine the three dimensional capabilities of the CIJ system, a spiral conductive trace was produced on the outer surface of a small glass cylinder (Figure 12).

The glass bottle was preheated for fast material evaporation so that undesirable trace flow was reduced. The final trace produced had uniform thickness, good electric conductance (resistance of and a strong adherence to the glass substrate. To further demonstrate the ability to print in three-dimensions, antenna patterns were printed on the inner surface of a bugle-shaped glass cup. To print the patterns, the cup was fixed on the 2-D motion table at a 45o angle to the axis of the inkjet head, as indicated by the dashed lines in figure 13(a). With precision modulation of the droplets and control of motion table, the printed traces on the three-dimensional curved surfaces were continuous and of consistent quality (figure 13b). Two very important characteristics of the printed system are demonstrated in traces shown in Figure 13b.

Two very important characteristics of the printed system are demonstrated in traces shown in Figure 13b. First, the system is capable of producing quality traces over a very long throw distance, as the length from the orifice to the bottom of the bugle shaped cup is about 4 inches. Secondly, the angle between the axis of orifice and the normal axis of deposited surface varies from -45 o to 45o, which demonstrates that the CIJ technique is capable of printing on complex 3-D surfaces.8 E −6.4ohmmeter)Summary:In an effort to reduce the cost of electronics fabrication, novel conductive ink materials are introduced in the present investigation that are capable of being printed in three-dimensions, using CIJ technology. Based on experiments performed using an inkjet printing system developed by the investigators, the following conclusions can be drawn:1.

Based on experiments performed using an inkjet printing system developed by the investigators, the following conclusions can be drawn:1. Specified aqueous silver ink solution can be inkjet printed to produce conductive metal patterns that have sufficient electric properties for use in commercial applications.2. The most important ingredient in the aqueous ink is silver compound, i.e. silver nitrate. Commercially available additives are required to make the silver nitrate a better conductive precursor. And the aqueous ink can be prepared instantly by simply mixing the ingredients with water. This makes the aqueous ink inexpensive, convenient, and compatible to the current CIJ technology.3433. Based on wear tests using a pin-on-disk tribometer, the ink produces conductive lines with impressive adherence to glass substrates. In addition, the traces produced had an electrical resistivity of only 2.9 times of bulk silver after curing.4.

In addition, the traces produced had an electrical resistivity of only 2.9 times of bulk silver after curing.4. Utilizing a cylindrical glass cylinder, the investigators showed that the new process was able to print the conductive traces in three-dimensional space. It is believed that further refining the CIJ process and ink material will allow the process to be used in the electronics fabrication industry.Figure 10 Stable droplets generated and charged from CIJ systemFigure 11 Sample silver heating gridFigure 12: Spiral Conductive TraceFigure 13 (a) The cup was fixed on the 2-D motion table; (b) Antenna pattern was printed on the 3-D inner surface of bugle cup with CIJ printing.344References1. M.Orme, J. Courter, Q.Liu, J.Zhu and R.Smith, “Charged molten metal droplet deposition as a direct write technology”, MRS 2000 Spring Meeting, SanFrancisco, April, 20002. Wallace W.Carr, Jeffrey F. Morris, F.Joseph Schork, Wayne C.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCONTINUOUS LASER SCAN STRATEGY FOR FASTER BUILD SPEEDS IN LASER POWDER BED FUSION SYSTEMH. Yeung1, B. Lane1, J. Fox1, F. Kim1, J. Heigel1, J. Neira21Engineering Laboratory, 2Physical Measurement Laboratory, National Institute of Standards and Technology, Gaithersburg, MD 20899AbstractResearch has shown significant influence of laser scan strategy on various part qualities in the laser powder bed fusion additive manufacturing process. The National Institute of Standards and Technology developed the Additive Manufacturing Metrology Testbed, which provides open architecture for flexible control and monitoring during a laser powder bed fusion additive manufacturing process. This allows extended control of scan strategies, including control of laser power and speed within each scan line.

This allows extended control of scan strategies, including control of laser power and speed within each scan line. A ‘continuous’ scan strategy can reduce build times and improve throughput by negating the need to turn the laser off between scan tracks (e.g., sky-writing). Also, less frequent laser power interruption can potentially improve the melt-pool continuity. Multiple experiments are performed utilizing the continuous and traditional scan strategies, and comparisons are made between build time and measured melt-pool qualities.IntroductionLaser powder bed fusion (LPBF) is an additive manufacturing (AM) process in which a focused, high power laser selectively melts geometric patterns into layers of metal powder, ultimately building a near fully dense freeform part.

The LPBF fabricated part quality is determined by many process parameters [1], such as the laser scan strategies (position, power, and velocity) and their respective synchronization, in conjunction with powder layer parameters (material, relative density, layer height, etc.). Varying the relative combination of these parameters can introduce known defects that plague LPBF parts. Pores, for example, have been attributed to various phenomena related to the power-velocity attributes or scan strategies (e.g., keyholing and collapse at high laser energy densities [2], or insufficient re-melting of adjacent scan vectors due to wide hatch spacing [3,4]). For example, Khairallah et al. noted that turning the laser off at the end of a scan vector can potentially cause pores to be trapped under the rapidly solidified melt pool, and recommended laser power decreased at these locations [5].

More adequately controlled velocity or power profiles along each scan vector can reduce probability of pore formation, or provide a parametric space for other property optimization. Apart from solidification physics at the end of a single vector, the general size, shape, and timing of a laser scanning raster pattern are known to affect the melt pool thermal history of the part, thus the resulting local and global residual stress and microstructure [6–8].Though LPBF technologies are rapidly improving and maturing, there is still wide potential for research in melt pool, scan track, and layer formation process physics. To fully define scan strategies, a system needs to control both how to fill up the build areas with scan vectors (hatching), and position, power, and velocity of individual scan vectors. The National Institute of Standards and Technology (NIST) designed and built an open architecture Additive Manufacturing Metrology Testbed (AMMT).

The National Institute of Standards and Technology (NIST) designed and built an open architecture Additive Manufacturing Metrology Testbed (AMMT). The control software developed for AMMT utilizes the full access of system parameters to enable quick deployment of different scan strategies. This work compares three scan strategies implemented on AMMT: continuous linear, continuous concentric, and constant build speed (sky-writing). ______________________________Certain commercial entities, equipment, or materials may be identified in this document in order to describe an experimental procedure or concept adequately.

______________________________Certain commercial entities, equipment, or materials may be identified in this document in order to describe an experimental procedure or concept adequately. Such identification is not intended to imply recommendation or endorsement by the National Institute of Standards and Technology, nor is it intended to imply that the entities, materials, or equipment are necessarily the best available for the purpose.1423AMMT Scan StrategiesThe AMMT is a facility at NIST which enables flexible and well-characterized monitoring and control of the laser powder bed fusion (LPBF) process [9,10]. AMMT control adapts an open platform design, using stereolithography (STL) files, G-code, and xy2-100 industrial standards for parts definition, path description, and galvanometer control, respectively. Figure 1 shows a generic software architecture for the AM system. The highlighted parts are modules implemented for AMMT.

Figure 1 shows a generic software architecture for the AM system. The highlighted parts are modules implemented for AMMT. The step 3 ‘G-code Generator’ converts a two-dimensional shape into scan paths described by G-code [11]. A standard G-code line, traditionally used for numerically controlled machines, describes a ‘move’ with destination point and speed. The LPBF-specific G-code version for NIST AMMT also includes laser power information. Step 4 ‘G-code Interpreter’ translates this G-code into xy2-100 commands per the predefined interpretation mode. Xy2-100 commands [9,10] are digital positions updated every 10 µs to control a laser unit and galvanometer scanner.

Xy2-100 commands [9,10] are digital positions updated every 10 µs to control a laser unit and galvanometer scanner. Therefore, the laser scan strategies are implemented by both ‘G-code Generator’ and ‘G-code Interpreter’.Figure 1: AMMT control designG-code GeneratorTo build a layer, a two-dimensional shape must be filled up with scan vectors (hatching) which are spaced densely enough to melt the powder into a joint solidified region. The hatching pattern can be linear or concentric as shown in Figure 2a. The hatching lines can be scanned in an interleaved manner, or sequentially with alternative power as shown Figure 2b. A larger scan area can also be divided into smaller sub-regions (islands) and scanned in a certain sequence (Figure 2c). The combinations of above form the hatching strategy.

A larger scan area can also be divided into smaller sub-regions (islands) and scanned in a certain sequence (Figure 2c). The combinations of above form the hatching strategy. G-code Generator fills a two-dimensional shape with scan vectors per the predefined hatching strategy, and generates G-code representing the scan vectors.1424Figure 2: NIST AMMT hatching strategies: (a) Linear and concentric hatching patterns. (b) Interleaved scan and alternative power scan. (c) Islands with different scan sequence.G-code InterpreterThe motion and power control of the laser spot within a move, as well as the transition between two sequential moves, is implemented in G-code Interpreter. Earlier work by the authors [9] proposed three laser path modes and three laser power modes. A summary is given below.a) Laser path mode1. Exact stop mode – motion stops exactly at the end of each move with maximum allowable deceleration.

A summary is given below.a) Laser path mode1. Exact stop mode – motion stops exactly at the end of each move with maximum allowable deceleration. If there is a subsequent move, motion will start immediately again with maximum allowable acceleration until it reaches the programmed speed, or until it needs to decelerate again. 2. Constant build speed mode – motion is kept at constant speed for the whole move, if the laser power is on for that move (i.e., a build move). In order to achieve that, extra moves may be added before or after the original move to speed up or slow down the motion, but laser power is kept off for the added moves.3. Continuous mode - two sequential moves of different velocities are connected by an arc to allow a smooth transition of velocity. Deviation from the designed path within the maximum tolerance is allowed.b) Laser power mode1. Constant power mode – power is kept constant at the programmed level during each build move, regardless of scan speed.

Constant power mode – power is kept constant at the programmed level during each build move, regardless of scan speed. If a subsequent move has a different power level, control sets the power to the new level.2. Constant density mode - power/speed ratio (power density) is kept at a predefined constant during each move. This constant is not necessarily the same for all moves.3. Thermal adjusted mode - power level is adjusted according to the predefined thermal properties of the building process.1425Experiment designDifferent scan strategies can be programmed on AMMT by the combination of hatching strategies, laser path modes, and laser power modes. Seven experiments were designed. Figure 3 shows the actual laser scan tracks from each experiment. All experiments were conducted on a ground stainless-steel plate (100 mm x 100 mm x 6.4 mm) without powder. The scan speed is 500 mm/s, power is 200-watt constant, and hatch spacing is 200 µm.

All experiments were conducted on a ground stainless-steel plate (100 mm x 100 mm x 6.4 mm) without powder. The scan speed is 500 mm/s, power is 200-watt constant, and hatch spacing is 200 µm. Wider spacing is used so individual tracks can be observed. Table 1 lists scan area (length x width) and scan strategy for each experiment. Experiment 1, 4, and 5 used constant build speed mode, also known as ‘sky-writing’ on some commercial LPBF systems, and are therefore marked as conventional and used as a base line for comparison. Others are marked as ‘AMMT’ in Table 1 as authors are not aware of any similar scan strategies on commercial systems, although this does not mean that similar concepts have not been explored before [12,13].Figure 3. Scans on stainless steel plate with 500 mm/s speed and 200-watt constant laser power. (a) Single square scan experiments 1-3. (b) Island scan experiments 4-7.Table 1.

Scans on stainless steel plate with 500 mm/s speed and 200-watt constant laser power. (a) Single square scan experiments 1-3. (b) Island scan experiments 4-7.Table 1. Experiment parametersExp # 1 2 3 4 5 6 7Length x Width (mm) 2.5 x 2.5 2.5 x 2.5 2.5 x 2.5 20 x 10 20 x 10 20 x 10 20 x 10Scan StrategiesConstant build speed Continuous linear Continuous concentric Constant build speed Constant build speed - island Continuous linear - island Continuous concentric - islandRemarks Conventional AMMT AMMT Conventional Conventional AMMT AMMT1426Melt-pool AnalysisA key signature characteristic in LPBF AM processes is the melt-pool geometry. One can observe the melt-pool to study the influence of laser control on the process.

One can observe the melt-pool to study the influence of laser control on the process. In this study, in-situ high-speed coaxial imaging is used to measure the melt-pool area, and post-process confocal laser scanning microscopy is used to measure the melt-pool height.Melt-pool width measurementTo measure melt-pool area, a high-speed camera was setup coaxially with the laser beam using a dichroic mirror, imaging lens, and filter. Emitted light from the melt-pool, which is filtered at 850 nm (40 nm bandwidth), is imaged on the camera sensor with nominal 1:1 magnification and 20 μm pixel size. The camera is set to 50 000 frames/s, 10 μs exposure time, 128 pixel x 128 pixel window, and 8-bit dynamic range (grayscale). The gray levels are used to relate to melt-pool dimensions [14]. Contours, representing isotherm lines, can be drawn on the raw melt-pool image to represent equal intensity (Figure 4).

The gray levels are used to relate to melt-pool dimensions [14]. Contours, representing isotherm lines, can be drawn on the raw melt-pool image to represent equal intensity (Figure 4). A contour with intensity of 50 digital levels was found to equate to the physical melt pool width based on the ex-situ measured scan track width via microscope inspection. This digital level contour is then used to infer the melt pool boundary from the high-speed images and to calculate melt pool dimensions and area.Figure 4. Melt-pool image analysis. (a) raw grayscale image. (b) processed image. Gray contour lines show different intensity levels (DL); red line shows melt-pool orientation.Melt-pool height measurementConfocal laser scanning microscopy enables the reconstruction of three-dimensional surfaces from a set of images obtained at different focal depths. A topography of the scan track such as Figure 5a can hence be created, with depth (height) represented by color.

A topography of the scan track such as Figure 5a can hence be created, with depth (height) represented by color. It is a very useful tool to visualize the effects of scan strategy on surface profile. From the topography, scan tracks are always bounded by valleys (Figure 5b), likely caused by surface tension of the molten metal. Once the two boundary valleys are identified, the melt-pool height can be defined as the peak between these two valleys (Figure 5c). Ideally the melt-pool width can also be determined by distance between the two valleys, but surface roughness or unsteady melt-pool behavior can sometimes create a local minimum (valley), too. Nevertheless, the local maximum is generally unique, making it a reliable approach for melt-pool height detection.1427Figure 5 Melt-pool width and height detection from confocal microscopy. (a) Topography of the melt track from Exp. 1. (b) 400 µm x 400 µm three-dimensional plot centered at position marked by white x.

(a) Topography of the melt track from Exp. 1. (b) 400 µm x 400 µm three-dimensional plot centered at position marked by white x. (3) Cross section profile along the red line on three-dimensional plot. (d) Profile along the blue line.Experiment results and discussionExperiments 1-7 were conducted using AMMT on a bare metal (stainless steel) plate; no powder was added. The microscopic images of the scanned regions were shown in Figure 3. The processes were in-situ imaged by the high-speed coaxial camera at 50 000 frame/s with 10 µs exposure time. The images were processed to obtain melt-pool area, with the results plotted in Figure 6. The scanned regions were then examined by confocal microscope. The surface profiles are plotted in Figure 7 and Figure 8. The surface profile of an un-scanned region on the metal plate is also shown in Figure 7 as a reference.Figure 6 shows the melt-pool area from the high-speed coaxial image analysis.

The surface profile of an un-scanned region on the metal plate is also shown in Figure 7 as a reference.Figure 6 shows the melt-pool area from the high-speed coaxial image analysis. Dark lines plot melt-pool areas for each individual frame, and red lines are the means for each experiment with one standard deviation. Table 2 lists mean, standard deviation, and build time specified for each experiment.Figure 6 Melt-pool area measurement from coaxial high-speed image analysis. (a) Single square scan experiments (Exp. 1-3). (b) Island scan experiments (Exp. 4-7).Table 2. Average melt-pool area, standard deviation, and build time for Exp. 1-7.Exp.

(a) Single square scan experiments (Exp. 1-3). (b) Island scan experiments (Exp. 4-7).Table 2. Average melt-pool area, standard deviation, and build time for Exp. 1-7.Exp. #Mean melt-pool area (µm2 )1 299002 333003 397004 417005 400006 42500Standard deviation (µm2 )440047008900470083008100Build time (s)0.0980.0760.0882.2583.6002.75814287 50300159003.356Based on the in-situ melt pool imaging, and ex-situ topography results, the following observations are made:1. For single square scans (Exp. 1-3), continuous linear mode is the most time efficient, while continuous concentric mode has the largest mean melt-pool area. The increasing melt-pool area for Exp. 3 (Figure 6) should be due to the residual heat building up as concentric circles are getting smaller.2. For multi-square island scans (Exp. 5-7), the trend is consistent with single square scans. The time for continuous linear mode (Exp.

For multi-square island scans (Exp. 5-7), the trend is consistent with single square scans. The time for continuous linear mode (Exp. 6) improved by 23% over constant build speed mode (Exp. 5), and average melt-pool size for continuous concentric mode (Exp. 7) improved by 25% over constant build speed mode (Exp. 5).3. Island scans (Exp. 4-7) yield a larger melt-pool area than their single square counterparts (Exp. 1-3). Longer scan time and more localized scan path helped build up the residual heat. 4. Comparing Exp. 4 with 5, island scan strategy does not seem to have any advantage for constant build speed mode, in terms of both building time and melt-pool size. The laser was turned on and off very frequently for Exp. 5 because of the shorter scan distance. That might have balanced out the localized heat effect of the island scan.5. The standard deviations of the melt-pool areas are the smallest for Exp. 1 and 4.

That might have balanced out the localized heat effect of the island scan.5. The standard deviations of the melt-pool areas are the smallest for Exp. 1 and 4. The power density is more consistent in constant build speed mode.Figure 7. Surface profiles for Exp. 1-3. (a) Topographies. (b) Histogram of height data with normal distribution curve fitted. X-axis is the height in µm, Y-axis is number of points. (c) Enlarged view of (b).Figure 7a shows the topographies for Exp. 1-3, and a reference region (an un-scanned area on the same metal plate). Figure 7b is the distribution of the height for each point on the topography. Figure 7c is an enlarged view of Figure 7b, and shows how the height data spread out around boundaries. The height profile for the reference fits very well into a normal distribution curve, which represents the surface1429roughness of the original metal plate.

The height profile for the reference fits very well into a normal distribution curve, which represents the surface1429roughness of the original metal plate. All experiments resulted in a ‘rougher’ surface as the height data spread wider (Figure 7b and Figure 7c), especially for Exp. 1 which goes beyond +/- 10 µm range. A rougher surface may end up a poorer part quality in a three-dimensional build as it prevents even distribution of powder, or it may even jam the recoater blade. To further visualize how different scan strategies affect the surface profile, the top left corners of each experiment were enlarged and compared in Figure 8.Figure 8. Surface profiles for Exp. 1-3. (a) Topographies. (b) Height profile along the red line. (c) height profile along the blue line.For constant build speed mode (Exp. 1), the laser power was turned on at the beginning and turned off at the end for each scan line.

(c) height profile along the blue line.For constant build speed mode (Exp. 1), the laser power was turned on at the beginning and turned off at the end for each scan line. That created bumps at the beginnings and holes at the ends, as indicated by the red and purple arrows in Figure 8a. Continuous linear mode (Exp. 2) reduces the laser on and off frequency, but the sharp turn between adjacent hatch lines creates complicated local thermodynamics. That still resulted in ‘valleys’ around turns, but at a smaller scale compared to Exp. 1. The continuous concentric mode (Exp. 3) has no sharp turns (except at the center), resulting in the most even surface, as shown in Figure 8b and Figure 8c.Summary and Future workThere is an open field of research into laser scan strategy with the potential to reduce defects, control residual stress or microstructure, or improve the speed and efficiency of material consolidation.

The NIST AMMT provides control and monitoring capabilities for such research. This paper demonstrated two advanced scan strategies implemented on AMMT – continuous linear and continuous concentric. The results show clear advantages over traditional (constant build speed) scan strategy for both efficiency and1430quality. The build time for the continuous linear strategy is 23% faster than the traditional strategy, and the continuous concentric strategy effectively reduced bumps and holes compared to the traditional strategy.Constant laser power was used for all experiments in this work, although AMMT has the capability for adjusting laser power at 100 kHz. A proper tuning of laser power should be able to further improve the surface quality. It is also shown here that a continuous island scan produced a larger melt-pool, and the island scan is believed to be able to reduce residual stress as well.

It is also shown here that a continuous island scan produced a larger melt-pool, and the island scan is believed to be able to reduce residual stress as well. A study of multilayer powder builds based on varying power continuous concentric island scans should be very interesting.References[1][2][3][4][5][6][7][8]Mani M, Lane B, Donmez M A, Feng S, Moylan S and Fesperman R 2015 Measurement science needs for real-time control of additive manufacturing powder bed fusion processes (Gaithersburg, MD: National Institute of Standards and Technology) King W E, Barth H D, Castillo V M, Gallegos G F, Gibbs J W, Hahn D E, Kamath C and Rubenchik A M 2014 Observation of keyhole-mode laser melting in laser powder-bed fusion additive manufacturing J. Mater. Process. Technol. 214 2915–25 Thijs L, Verhaeghe F, Craeghs T, Humbeeck J V and Kruth J-P 2010 A study of the microstructural evolution during selective laser melting of Ti–6Al–4V Acta Mater.

How- ever, modeling material gradients is still tedious with current approaches, especially when smooth transitions are required. Current approaches can be distinguished into a) NURBS-BReps-based and b) voxel-based. In case of NURBS-BReps, discrete material distributions can be modeled by manually introducing separate shells inside the object; smooth gradation can only be approx- imated in discrete steps. For voxel representations, gradation is discrete by design and comes along with an approximation error. In addition, interacting on a per-voxel basis is tedious for the designer/engineer. We present a novel approach for representing material gradients in volumetric models using subdivision schemes, supporting continuity and providing elegant ways for interac- tive modeling of locally varying properties.

Additionally, the continuous volumetric representation allows for on-demand sampling at any resolution required by the 3D printer.IntroductionModern 3D printers can vary mechanical properties of objects locally, e.g. by mixing dif- ferent base materials or varying process parameters – the latter is less common, still. Traditionally,1675CAD systems using BRep-NURBS only allow / support assigning a material for each solid or topo- logical lump. However, modern 3D printers, such as the Stratasys J750, allow for material varia- tion on a per-voxel basis. Obviously, deﬁning a lump per voxel would be an impractical approach considering the size of the individual “elements” being printed. Of course, one could consider coarsening the voxels representing the geometry of the object, but that would limit the accuracy of the shape representation and the freedom in material variation.

Of course, one could consider coarsening the voxels representing the geometry of the object, but that would limit the accuracy of the shape representation and the freedom in material variation. Alternatively, it is possible to assign material information on a per-slice level, but slicing is an approximating discretization performed in a post-process to the modeling stage. Thus, the actual geometry would not carry the material information. Considering the huge number of slices created, assigning material information would again be a tedious process.So, wouldn’t it be ideal to enable the designer to deﬁne material properties on a contin- uous representation allowing for editing on a moderate number of control points? For geometry, CAD systems use splines to represent continuously differentiable shapes.

In isogeometric analysis (IGA), the same concept is extended towards modeling and simulating on the same representation scheme (trivariate splines) and has recently been used to assign materials to volumes. However, trimmed splines are not necessarily the most easy-to-use representation for organic shapes as can be seen in the movie and game industry where subdivision schemes are much more common than in mechanical engineering.In this paper, we propose subdivision schemes to model shape and material property dis- tribution for 3D-printed objects, because 3D printing is particularly well suited for generating organic (bionic) shapes that cannot be manufactured with traditional subtractive production tech- nology. Leveraging subdivision allows us to enable the designer to deﬁne the shape and properties he or she wants to achieve at certain regions in space on the same representation scheme.

Leveraging subdivision allows us to enable the designer to deﬁne the shape and properties he or she wants to achieve at certain regions in space on the same representation scheme. The con- tinuous representation can be evaluated at any point in space with the resolution of the 3D printing machine – no discretization or conversion required. The properties deﬁned on a control mesh are smoothly interpolated into the interior of the object. The interpolated property is matched as closely as possibly by mixing the materials provided by the 3D printer according to their properties / behavior. Properties can also be assigned to internal control points giving the needed freedom to the user to specify continuously graded material properties throughout the object.Related WorkModeling for Additive Manufacturing: In engineering, the shape of a part is usually de- ﬁned using computer aided design (CAD).

For this purpose, trimmed NURBS patches, which bound the part to be produced as a continuous surface, have been established in industrial practice as the standard representation for the description of free-form geometry. Subdivision surfaces gen- erated by methods such as Catmull-Clark [1] or Loop subdivision [2] are an alternative modeling technique for geometric objects. Algorithms for efﬁcient and accurate representation of subdi- vision surfaces [3] and methods for sketch-based generation of subdivision models [4, 5] have already been developed. These forms of surface representations are now established as a standard in computer graphics and are gaining increasing acceptance in industrial surface design. However, although closed surface meshes implicitly describe a volumetric object, they are not considered to be volumetric meshes, since the do not contain any information about the interior of the volume1676they enclose.

Common tools for polygonal modeling used in design and computer animation, such as Maya [6], Blender [7], Rhino [8], and SolidWorks [9] only support surface modeling.For fully volumetric models, different representations have to be used. In the context of IGA [10], trivariate NURBS volumes are used as a representation for volumetric models. In 2016, Massarwi and Elber presented a framework for modeling volumetric B-spline models in their so- called “VRep” (“Volume Representation” analogously to the commonly used “BRep”) [11]. Vol- umetric subdivision methods, such as the ones presented by Joy and MacCracken [12], Schaefer et al. [13], or Chang et al. [14], transfer the concept of subdivision surfaces to volumetric meshes. In addition to vertices, edges, and faces, a volumetric subdivision mesh consists of cells and has inner faces to separate them.When it comes to assigning materials to volumetric models, two main concepts exist: image-based and voxel-based.

For image-based material assignment, a 2D image is created for every layer, carrying the material values in its color channels. With voxel-based material mod- eling, material values are explicitly assigned to individual cells in a 3D voxel grid. GrabCAD Print [15] and Autodesk Monolith [16] are just two tools on the market that allow volumetric ma- terial editing with the above techniques. However, for both concepts, material gets assigned to the model after slicing/voxelization. In contrast, our approach operates on the original model before slicing/voxelization takes place and is therefore independent from the ﬁnal printing resolution and orientation.Exchange Formats for Volumetric Models: Currently, the primary exchange format for 3D-printed models is the stereolithography (STL) ﬁle. Standard STL ﬁles only contain a set of triangles describing the outer surface of a mesh, without any information about color, material or scale.

Standard STL ﬁles only contain a set of triangles describing the outer surface of a mesh, without any information about color, material or scale. While more advanced formats exist, they all describe the model only by its boundary, i.e., as a boundary representation. For example, the 3D Manufacturing Format (3MF), developed by a consortium large CAD, CAM, and 3D-printing companies, adds colors, thumbnail images, and digital signatures but is still based on triangulated surface meshes (see, e.g., the 3MF speciﬁcation [17]).While these advanced formats allow assigning multiple surface colors via vertex colors or even using texture images, actual material types can only be speciﬁed by name, turning mapping to an actual material into a manual process:<base> attribute name: Speciﬁes the material name, intended to convey design intent, for the purpose of aiding users in mapping to print materials.

[17]Furthermore, due to the BRep nature of current formats, material gradations or mixes within the volume cannot be described directly, even though current 3D printers can already produce material mixes. Approximations such as multiple thin shells are used and often created using a tedious, manual process.1677ModelingAs mentioned earlier, we investigate the use of subdivision schemes to model shape and ma- terial property distribution. Subdivision models create continuous shapes while supporting com- plex and unstructured topologies.(a) Freeform(b) Block-based(c) Sketch-basedFigure 2: Three subdivision models created with different modeling techniques. (a) has been created with freeform polygonal modeling, while (b) has been created with a block-based approach and (c) is the result of sketching.Subdivision Modeling: Different techniques exist to create subdivision models.

Although most tools implement polygonal modeling (block-based modeling), procedural modeling with an- alytical functions as well as sketch-based modeling are also possible. Sketch-based approaches are particularly useful for modeling strut-heavy structures that are most often the result of topology optimization techniques. Figure 2 shows three subdivision models created with different modeling techniques.Material Deﬁnition using Subdivision Models: In addition to the geometric modeling, the modeling of material distributions constitutes an important extension for 3D printing. When applying the subdivision rules to a given control mesh, a reﬁned set of points is created by interpo- lating the original ones. When representing geometry, every control point only has a 3D position (x, y, z). However, subdivision rules and the underlying interpolation scheme can be applied to points of arbitrary dimension. Therefore, we introduce higher-dimensional control points, e.g.

However, subdivision rules and the underlying interpolation scheme can be applied to points of arbitrary dimension. Therefore, we introduce higher-dimensional control points, e.g. 4D, 5D or 6D, with a geometric position in 3D space as well as material parameters such as density, Young’s modulus, or Poisson’s ratio. Figure 3 shows an example of control points with attached stiffness values. Note how the stiffness values are interpolated over the subdivided geometry, cre- ating a smooth gradient. Following the subdivision scheme, material properties can also be deﬁned volumetrically, when creating models with inner control points. This way, models with a soft outer shell over stiff internal structures can be created and very interesting physical behaviors can be achieved.Subdivision schemes – both for surfaces, as well as volumes – converge towards a smooth limit surface/volume when applying a theoretically inﬁnite number of subdivision steps.

For most schemes, the limit can be evaluated mathematically, without performing explicit subdivision steps at all. By “attaching” additional parameters directly to the control points, the limit values of these parameters can be evaluated everywhere in the volume with the same algorithm that is used for1678(a) Subdivision(b) Limit(c) Inner control pointFigure 3: Material properties assigned at the control points and their effect on the subdivided model and the limit. (a) shows an example of a subdivision cylinder, showing the effects on the subdivided surface, while (b) shows the corresponding limit surface. (c) shows a cut through the object, showing inner control points. In this example, white color represents a low stiffness, while blue color represents a high one.evaluating the limit positions.

(c) shows a cut through the object, showing inner control points. In this example, white color represents a low stiffness, while blue color represents a high one.evaluating the limit positions. Figure 3(b) shows an example of evaluating stiffness values on the limit surface as well as in the limit volume.Interaction Techniques: For evaluating our approach, we implemented a prototype appli- cation. It is written as a C++ desktop application and provides two different interaction techniques for applying material properties on the control polygon of a given subdivision model. One way is to directly select individual control points to allow a targeted, local modiﬁcation. The other way is to use brushing to quickly and intuitively edit a larger number of nodes that are within the brush’s radius. Figure 4 shows examples of both techniques.

assigned to one control point. Finally, (c) shows a brushing operation, changing the material parameters for multiple control points at once. Blue color corresponds to very stiff and white color to a very ﬂexible material parameters.1679Export and Storage of Volumetric Multi-Material ModelsNow that our subdivision model has been created and the distribution of its material param- eters has been deﬁned, the model has to be prepared for 3D printing. As available slicers cannot process subdivision geometry directly, but require tessellated input, we discretize our model into a tetrahedral mesh and store it using a volumetric attribute-based data format.Implicit Tetrahedrization: As presented in our previous paper [18], we exploit the already existing volumetric structure given by the subdivision scheme to create the tetrahedral cells out of the hexahedral subdivision cells resulting from n explicit subdivision steps.

The number of subdivision steps n can be increased or decreased in order to control the resolution of the resulting tetrahedral mesh. As the entire meshing process can be expressed as a set of pre-computable matrix-matrix multiplications and a ﬁnal matrix-vector multiplication, the tetrahedral mesh can be re-generated very efﬁciently, when modifying the geometry or the material parameters on the original control mesh – as long as the topology of the control mesh remains unchanged. All the calculations to get from the set of original control points to the tetrahedral mesh are performed independently on the x, y and z coordinate of every three-dimensional control point. Therefore, it seems natural to apply the same calculations also on the material properties when dealing with 4D, 5D or 6D control points described in Section 4.2.

Therefore, it seems natural to apply the same calculations also on the material properties when dealing with 4D, 5D or 6D control points described in Section 4.2. The material parameters attached to every control point are “subdivided” and “meshed” in the same way as the control point’s geometric position.Export and Storage Format: Since current exchange formats cannot represent volumet- rically varying material parameters efﬁciently, we propose a volumetric, attribute-based model Instead of describing a volume’s boundary using a triangle mesh, we use a exchange format. tetrahedral mesh to describe it’s volume. Furthermore, instead of only storing material colors and names, we also store physical properties such as density, Young’s modulus and Poisson’s ratio. It is then up to the printer driver to approximate these properties as well as possible using techniques such as material mixes (potentially via dithering) or metamaterials (mesostructures).

It is then up to the printer driver to approximate these properties as well as possible using techniques such as material mixes (potentially via dithering) or metamaterials (mesostructures). This mirrors how 2D color printers receive a color speciﬁcation which they approximate using halftoning or contoning, without the user having to select speciﬁc inks or dot patterns. Furthermore, by storing physical properties with the model, it becomes directly simulatable.The storage format is based on HDF5 2.0 [19], a self-describing, hierarchical, binary ﬁle-format that supports n-dimensional datasets, stores metadata, has free cross-platform imple- mentations, and supports compression. Extensibility is achieved through the use of universally unique identiﬁers (UUID) [20], 128-bit identiﬁers out of which 122 are random for approximately 5.3 · 1036 possibilities. As this effectively rules out collisions, no central administration of exten- sions is required.

As this effectively rules out collisions, no central administration of exten- sions is required. When adding a new physical property, a new UUID is chosen and used from that point on to describe that property’s semantic meaning. Additionally, we store a user-readable name and SI-unit exponents to clarify intent if the UUID is not yet known by the tool used.To be able to describe sharp edges as well as gradients, we use a multi-indexed approach. Each of the four vertices of a tetrahedron is given an index into the position property that also describes the model’s topology. Additional indices are attached to each of these four indices for1680every physical property.

Additional indices are attached to each of these four indices for1680every physical property. Therefore, neighboring tetrahedra can use the same position but different property indices to create a sharp transition.We explicitly do not store any slicing information, resolution or speciﬁc materials, as all ofthese would closely couple a model to a speciﬁc printer model, vendor, or technology.Volumetric SlicingTo print the 3D objects with a 3D printer that supports multiple materials, the objects have to be converted into the according input format for that particular printer. In case of the Stratasys J750 printer that we used for the experiments in this paper, as well as for other printers, the input format is a stack of 2D images that determine, which material is placed at which position in a three-dimensional voxel grid. Every image represents one layer. These layers are ﬁnally printed on top of each other.

Every image represents one layer. These layers are ﬁnally printed on top of each other. Creating the layer information from the input 3D object is called “slicing”.The slicing process is performed by the software Cuttleﬁsh [21] and consists of two steps: First the surface of the exported tetrahedral mesh is rasterized and after that the voxelization of the interior takes place.Surface rasterization: During the surface rasterization step, the input 3D object’s surface is tested against the three dimensional voxel grid and all voxels that intersect with the surface of the object are marked as occupied surface voxels. The implemented surface rasterization method is described in Section 3.4 of [22].Interior voxelization: The interior of the 3D object is then ﬁlled by evaluating the winding number of the voxels as described in [23]. To speed up the process, a subset of all voxels is selected as seed points, for which the winding number gets evaluated.

To speed up the process, a subset of all voxels is selected as seed points, for which the winding number gets evaluated. The information of these seed points is then propagated throughout the volume by using a ﬂoodﬁll approach. An overview of this method can be found in Paragraph 108-110 of the patent [24].After the voxelization process, the application has performed an inside-outside segmenta- tion of the set of voxels and now materials must be assigned to create the output for the 3D printer. For this we interpret the material values per vertex of every tetrahedron as tonal values – prob- abilities that a material is placed. These attributes are assigned to the surface voxels during the rasterization phase by interpolating between the three vertices that form a face of the tetrahedron.

These attributes are assigned to the surface voxels during the rasterization phase by interpolating between the three vertices that form a face of the tetrahedron. As described in Section 4 of [25], the material values are then interpolated from the surface into the volume of a voxelized tetrahedron which results in the assignment of tonal values to each voxel. These tonal values are then halftoned using a 3D blue noise mask, which results in a concrete assignment of materials.The described implementation is missing a linearization step that takes the per vertex color, interprets it as stiffness, and converts it into a material probability that correspond to that stiffness. One possible approach to obtain a linearization is to print targets with different material mixtures and measure their stiffness.

One possible approach to obtain a linearization is to print targets with different material mixtures and measure their stiffness. The resulting data set will plot the non-linearity of the stiffness and interpolating between the samples will allow us to obtain a material probability for stiffness that can1681be achieved with this material. In case more materials are used, the sampling dimension increases and the interpolations have to be performed in higher dimensions. This described missing workﬂow is implemented in Cuttleﬁsh although sample data is missing.ResultsTo print our example models, we used a Stratasys J750 printer and a variety of photopoly- mer materials. As a stiff material we decided on a cyan Vero photopolymer. As a ﬂexible material we chose Agilus30, a PolyJet photopolymer with high tear-resistance, capable of withstanding repeated ﬂexing and bending.

As a ﬂexible material we chose Agilus30, a PolyJet photopolymer with high tear-resistance, capable of withstanding repeated ﬂexing and bending. To represent our smooth gradation of material properties, both ma- terials are blended together with the corresponding ratio for every voxel. Some of the printed parts alongside their multi-material digital counterparts are shown in Figure 5. The printing process for a single model takes approximately 1 hour and 45 minutes and uses about 110 grams of material. The exact numbers depend on the actual material distribution.(a) Printed parts(b) ModelsFigure 5: Printouts of multiple example models with the Stratasys J750 (a) and the corresponding digital models created with our approach (b). Blue areas indicate a high stiffness, while white areas indicate a low one.Conclusion and Future WorkIn this paper, we present a novel approach for deﬁning volumetrically varying material parameters in 3D models.

By attaching properties such as density, Young’s modulus, or Poisson’s ratio to the control points of a subdivision control mesh, we can represent smooth gradients of parameters across and inside the model when applying explicit subdivision steps or direct limit evaluation. The approach offers great ﬂexibility in geometric modeling and property deﬁnition with a relatively low number of degrees of freedom. The ability to directly work with a volumetric representation enables great exploitation of modern AM possibilities.With our approach, we deﬁne smooth variations of material properties. However, even though the chosen subdivision scheme supports crease edges and sharp features for the geometry, discrete boundaries and discontinuous transitions are not possible for material properties.1682For Catmull-Clark subdivision models, the property gradation follow higher-order basis functions similar to the ones of cubic B-splines.

When converting the subdivision model into a tetrahedral mesh for slicing, we discretize the smooth material properties for every tetrahedron. Instead of calculating constant properties for every element by averaging the values at its corners, the value at a given position inside the tetrahedron could be approximated using the same subdivi- sion basis functions as we do for the modeling. To speed up the subdivision and tetrahedrization process, we plan a parallel GPU implementation using a highly efﬁcient GPU data structure [26]. In order to eliminate the tetrahedrization completely, direct slicing of the subdivision geometry is envisaged.AcknowledgementsThis work was supported by the EU-project CAxMan (project number 680448) which is co- funded by the Horizon 2020 Framework Program H2020-FoF-2015 of the European Union. More information can be found at https://www.caxman.eu/.

More information can be found at https://www.caxman.eu/. We thank Alan Brunton for his implemen- tations of the voxelizer and the distance transform as well as André Schulz for the implementation of the halftoner.References[1] E. Catmull and J. Clark, “Recursively generated B-spline surfaces on arbitrary topologicalmeshes,” Computer-Aided Design, vol. 10, no. 6, pp. 350–355, 1978.[2] C. Loop, “Smooth subdivision surfaces based on triangles,” Ph.D. dissertation, University ofUtah, 1987.[3] V. Settgast, K. Müller, C. Fünfzig, and D. Fellner, “Adaptive tesselation of subdivision sur-faces,” Computers & Graphics, vol. 28, no. 1, pp. 73–78, 2004.[4] M. Bein, S. Havemann, A. Stork, and D. Fellner, “Sketching subdivision surfaces,” in Pro- ceedings of the 6th Eurographics Symposium on Sketch-Based Interfaces and Modeling. ACM, 2009, pp. 61–68.[5] Z. Ji, L. Liu, and Y.

ntinuous Sculpting as a Novel Method to Reduce the Staircase Effect in Solidfreeform Fabrication Processes that are Dependent on Phase Transformation.Rajeev Dwivedi*, Indira Dwivedi†, Bharat Dwivedi†STEM and Robotics Academy, Plano TX †Lake Washington School District, Sammamish WAAbstractInherent to 2-1/2 axis deposition in Solid Freeform Fabrication is Staircase effect. The surface geometry of parts produced using Solid Freeform Fabrication, therefore is not smooth. Continuous sculpting is a novel method that combines additive manufacturing to surface modification by relative motion of a forming tool. Molten material is added to substrate. A forming tool that can tilt along different axes is connected to the end effector. The forming tool is oriented along a desired angle and then applies force on the molten material to give it desired shape on the fly.

The forming tool is oriented along a desired angle and then applies force on the molten material to give it desired shape on the fly. This paper introduces the novel method and a simple case studies.IntroductionPast few decades have seen significant development in the field of solid free form fabrication. The developments have been done in field of processes, materials, and computer representation of the parts. However, inherent to the 3D printing is 2-/12 axis-based deposition that make parts susceptible to staircase effect[1,2]. Parts produced by such processes are limited in meeting functional intent or undergo significant post processing before they can be used. Parts manufactured by Solid Freeform Fabrication as it cannot be used for systems and assemblies that require sub millimeter order tolerances and part accuracy. Typically, the parts produced are primarily for the visualization and form fit evaluation or short duration functions.

Typically, the parts produced are primarily for the visualization and form fit evaluation or short duration functions. When features are critical, the staircase effects and a mismatch between the inter-layer thickness allowed by the SFF may lead to building the features partially. While such a problem may to certain extent be addressed by adaptive slicing; for the high-end parts especially the ones that are finish and feature form critical, may require significant pre- and post-processing.Many multi-axis CNC systems, Robotics systems as well as hybrid CNC-Robot systems have been used to add material, have been proposed [3,4,5]. Alternatively, a hybrid additive-subtractive processes have been proposed for producing high quality useable parts [6]. The systems are, however, extremely complex and require operator subject matter expertise.Most of the 3D printing platforms are based on a Cartesian architecture.

The systems are, however, extremely complex and require operator subject matter expertise.Most of the 3D printing platforms are based on a Cartesian architecture. Where displacement of end effector is performed along three cartesian axes and the material is extruded/added along one of the major axes. Usually, the axis of material addition is along the Z axis. Minimum resolutions within the form is dependent on stepper motor resolution and the linear actuator. Using the 3D printing, it is nearly impossible to produce a high-end functional part by only additive process.

Using the 3D printing, it is nearly impossible to produce a high-end functional part by only additive process. To meet the dimensional and material specifications, the subtractive processes must go hand in hand with the additive process.1382Many other SFF dependent processes [7] such as development of scaffolds for use in cell-based therapies to repair damaged bone tissue cannot use subtractive post processing for form fit improvements therefore, the additive process must in first pass manufacturing minimize the staircase effects. The residue created from the scaffold may impact the bone tissue generation.This paper describes an alternate approach being explored to reduce the staircase effect by mechanically forming semi-solid material before it is solidified. The paper describes the architecture of the test system and specifics of the equipment used.

The paper describes the architecture of the test system and specifics of the equipment used. The paper also describes the preliminary findings and suggests a path forward to implementation.Additive-subtractive process integration The 2-1/2 axis deposition mode assumes that the material addition is such that each layer is deposited to enable a uniformly thick layer. Between two subsequent layers, geometrically there is no variation along the axial direction. As described in the Figure 1, any critical features lying within twoFigure 1: An example of functionally critical feature missed due to staircase effectadjacent layers will be missed.Another issue worth noticing is that at the fundamental level, the material is dispensed in non-square profile (Figure 1). This may be attributed to the surface forces/surface tension of the fused liquid.

This may be attributed to the surface forces/surface tension of the fused liquid. The uniformity of the layer surface profile is aggregate of material dispense along different regions of aFigure 2 : Non-square profile along the outer surface of 2-1/2 axis depositionlayer. Remelting of previously deposited material and mixing with the newly added material enhances the surface uniformity; however, there are many factors that limit the quality of the surface. The energy may not be sufficient to melt the previously deposited material adequately. The material properties may not be amenable to frequent remelting. Relying purely on the material fusion to ensure a flat surface makes it extremely difficult to quantify the extent of variation.

The material properties may not be amenable to frequent remelting. Relying purely on the material fusion to ensure a flat surface makes it extremely difficult to quantify the extent of variation. For the additive processes the1383amount of material added depends on (1) the rate of material dispensing and (2) the relative speed of the additive end effector and the substrate.Many SFF processes resort to the combination of additive and subtractive manufacturing operations to build superior quality functional parts. Material deposition and intermittent machining to (1) restore a datum for critical features (2) eliminate significant geometric variations (3) Add blind features (4) remove sacrificial material needed to support overhanging regions. While many researchers have frequently used a complex architecture that is based off two spatial manipulators one additive and one subtractive, the same would not be advisable for the following reasons-1.

Multiple manipulators would need to establish corresponding multiple datum points. For a high quality part, a range of accessory implements and feedback transponders will be needed to minimize the system to system variation.2. Two separate set of manufacturing data needs to be generated. While CNC code can incorporate not only linear but also many curvilinear, path segments the same for 3D printing is based on linear segments. The STL formatted file and subsequent manufacturing data produced from the STL file comprises of linear segments. It is possible to utilize the high resolution path traversal and therefore superior geometric attributes with the CNC codes,Forming tool to reduce the staircase effect by continuous casting Concept proposed in this paper suggests using an actuator to sculpt the deposited material before it is fully solidified. Figure 3 describes the process steps for continuous sculpting. A plunger is placed within proximity of the material deposition head.

Figure 3 describes the process steps for continuous sculpting. A plunger is placed within proximity of the material deposition head. After a new layer of material is deposited the plunger is actuated. The plunger pushes the molten material to imprint form. Next the plunger is retracted, and the material is allowed to solidify.Figure 3 : Steps for continuous sculpting as a method to reduce staircase effect.As described in the Figure 4, the angle of the plunger may be adjusted per the desired surface profile.1384Figure 4 : Pivot and plunge of the tool to enable desired surface profileThe material delivery by extrusion and the path followed is a continuous process however, the plunging is done intermittently. The plunging action is required to separate the tool from molten material and prevent deformation due to adhesion of melted material onto the tool. The frequency of the plunging tool must enable overlap such the surface profile is smooth.

The frequency of the plunging tool must enable overlap such the surface profile is smooth. As shown in Figure 5, if the plunger frequency is inadequate and does not allow overlap in the indentations the surface quality is impacted.Figure 5: The impact of plunger frequency on the roughness of the surface profilePrototype experimental equipment1385Figure 6 : Section view of modified end effectorThe experimental setup is created by retrofitting a Creality Ender 3, PLA compatible 3D printer. As described in Figure 6, a shell is attached to the extruder with a hole to enable smooth passage of 0.5mm diameter stainless steel wire. The hole diameter for passing the wire is 0.6mm and the hole is drilled at 55degrees. Due to limitation of the available hardware the current setup has a fixed plunging angle.The plunger hole is 0.25mm offset from the plane of extrusion. Allowed speed limit of printing for the end effector is 60-100mm/s.

Allowed speed limit of printing for the end effector is 60-100mm/s. However, the experiments are performed at print speed of 75mm/sec. The extruded material is 1.75mm PLA.Figure 7 describes the actuation mechanism. The actuation mechanism is built using a linear screw drive. The linear drive is activated by a 2-phase 4-wire 22.5 Degree Stepping Angle Stepper Motor FSM0815-KD95. The linear screw drive is based on a Screw Diameter 2.5mm and Screw Length of 43mm. For actuating the mechanism, itself a general purpose open source microcontroller, Arduino UNO is used. The driver for the motor is a L298 Stepper driver. During the deposition actual plunge frequency is 1/sec.Figure 7: Actuation mechanismInitial results and findings1386Figure 8 : Results from 1 plunge/secFigure 8 describes the surface smoothening by plunger actuation. The stepper motor-based mechanism limited the plunge frequency to create a smooth uniform surface.

The stepper motor-based mechanism limited the plunge frequency to create a smooth uniform surface. The 3D cad model of the part is created with surface over-compensated for the staircase effect.Figure 9 : measured values across 4 layersFigure 9 shows measured values for 4 layers where the desired layer height of approximately 0.25mm, a compensation of 0.125mm was done for each layer. After the plunging the profile of the surfaces. As manufactured surface profile exhibits additional 0.125mm; however, the points along where the surface modification is done using the plunger the surface is in agreement with the desired surface profile. . Conclusion/Summary/ Future workA plunger-based continuous casting is done to improve surface profile of deposited layers. Preliminary results are described in the paper and exhibit improvement using continuous casting. Additional experiments and measurements are required to prove effectiveness of the approach.

Preliminary results are described in the paper and exhibit improvement using continuous casting. Additional experiments and measurements are required to prove effectiveness of the approach. Other non-sticking material and improved plunger tool will be used to prevent adhesion of molten plastic and the tool. Plunger will be developed further to add tilting such that the tool can adapt to local surface inclination.Resources:13871.2.3.4.5.6.7.Armani Amir, Li Wenbin, Leu, Ming, Landers Robert, Planning freeform extrusion fabrication processes with consideration of horizontal staircase effect. Solid Freeform Fabrication Symposium, 2015, Austin, TX, USAArni, R., and Gupta, S. K. (September 1, 1999). "Manufacturability Analysis of Flatness Tolerances in Solid Freeform Fabrication ." ASME. J. Mech. Des. March 2001; 123(1): 148–156.

ntour Following for Scanning Control in SFF Application : Control Trajectory PlanningYing-Jeng Engin Wu and J. J. Beaman The University of Texas at AustinAbstractGeometric contour following for scanning control in SFF application is used to refine the boundary ofthe parts for increasing the accuracy or to develop the capability to arrange various scanning directions and paths for improving the part strength. The scanners must be driven to follow the prescribed path as fast as possible, limited by available torques. In this paper the minimum time optimal control problem with specified path and limited control torque is formulated. According to the trade-off between various requirements, a control strategy is studied.IntroductionGeometric contour following for scanning control in SFF is used to refine the boundary of the parts for inqreasing the accuracy or to develop the capability to arrange various scanning directions and paths for improving the part strength.

The raster mode and vector mode in scanner application are common in industry. However, the former is not appropriate to this application,the latter is too<slow. Thus, it is important to develop a control strategy such that the scanners are driven to follow the prescribed path as fast as possible without exceeding the available torques. In this paper, the minimum time optimal control with specified path and limited control torque is formulated. One special point in SFF application is that the control torques and specified path are in different spaces shown in fig#l, like a two-degree-of..freedom rotational manipulator.For general geometric path, it could be defined by a parameter. Specially, arc length , s ' is used in this paper such that its fJISt and second derivatives present tangential velocity and acceleration.

Specially, arc length , s ' is used in this paper such that its fJISt and second derivatives present tangential velocity and acceleration. Since the path following problem is with one degree of freedom nature, by choosing the state variables as sand ds/dt, the problem could be converted to s-space, and phase plane technique is usable. Based on the Pontryagin's minimum principle, it is shown that there exists no singular subarc in the minimum time solution [1]. The same result is obtained and used by Bobrow, Dubowsky and Gibson ( 1985 ) [2], and Shin and McKay ( 1985 ) [3], when they develop their construction algorithms for switching curve in phase plane, which are adopted here because they are conceptually easy.The steady state solution for a circle path is obtained and compared with the maximum constant speed solution. Then, the transient parts are added.

Then, the transient parts are added. From the observations, a control strategy is proposed with a trade-off between various requirements as below:(l) The speed is as fast as possible with available bounded torque. (2) The speed is constantwise. (3) The speed is limited above by available Laser power. (4) The speed is limited below by the continuity of sintering which does not occur ifLaser power is too low.126Minimum Time Optimal Control ProblemGalvanometer systemLasery8x = G1 (x, y) 8v =G2( x, y)xWorking surfaceflg#l. Scanner SystemA. 8x-8y space: The state variables are chosen as 8x, cox, 8y, COy.tc min J = min f dt = min tc ° subjected to: d8x/dt = COx,dcox/dt = gl ( 8x, cox) + 'tx, d8y/dt = COy, dcoy/dt = g2( 8y, COy ) + 'ty,andM ~ 'ty ~ M, where ¥ is a positive valueM ~ 'tx ~ M, S(x,y) = 0, which is a simple closed curve with no sharp comer.

x(O) = xo, x(tc) = xc, vx(O) = VxO, vx(tC) = VxC, yeO) = YO, y(tc) = YC, Vy(O) = VyO, Vy(tc) = VyC, are given.with the nonlinear geometric relation between control space and specified path space as shown in fig#l :8x = Ol( x, Y) 8y = 02( x, y)then8x(0) = 8xo, 8x(tc) = 8xc, 8y(0) = 8yO, 8y(tc) = 8yC, are somefunctiofls of xo, xc, vxO, Vxc, YO, Yc, VyO, vyC·cox(O) = coxO, cox(tc) = COxC, coy(O) = (OyO, COy(tc) = COyc,Note that (01,02) is one-to-one mapping between (x,y) and (8x,8y) in the SFF application, and all the derivatives are continuous in working domain [6].127B. s - space : S(x,y) =i ocould be defined as x = f1 ( S ) and y = f2( s ), where s is the unit arc length, and the ftrSt derivatives of f1 ( s ) and f2( s ) are continuous since no sharp comer in the path.

Let v = ds/dt, u = d2s/dt2, 9x=Gl(f1(S),f2(S»=h1(S) 9y = G2(f1( s), f2(S » =h2( s ) rox = h1'( s) v, roy = h2'( s ) v,thendrox/dt = h1"( s) v2 + h1'( s) u, d<oy/dt = h2"( s ) v2 + h2'( s ) u,andNote that h1'( s ) and h2'( s ) are continuous [7].The minimum time optimal control problem may be converted astf min J = min f dt = min tf °subjected tods/dt = v, dv/dt = u, -M S; h1"( S ) v2 + h1'( s ) u - g1( h1( S ), h l'( S ) v ) S; M (C1) , -M S; h2"( s ) v2 + h2'( s ) u - g2( h2( s ), h2'( s ) v) S; M (C2)withs(O) = so, v(O) = yo,S(tf) = Sf, V(tf) = Vf,andBased on the Pontryagin's minimum principle, it is shown that there exists no singular subarc in the minimum time solution, the solution is always on one boundary of those inequality constraints.Solution of the Minimum Time Optimal ProblemFrom the Pontryagin's minimum principle, the two-point-boundary-value problem is obtained by the necessary conditions and solved by Shooting method [4] [5].

However, the difficulty is found after the solution of first example is obtained. There exist the second derivatives in those inequality constraints, which are put into the Hamiltonian. Then, the dynamic equations of Lagrangian multipliers contain the third derivatives, dlJdt = - aHlas, and the forth derivatives appear in th~ formula of shooting method. The numerical singularity and sensitivity of convergence must be handled. Based on the motivation as below, the phase plane technique is followed [2],[3].tr °Sfmin J = min f dt = min f (lIv)dssOIn the phase plane, minimizing the cost is conceptually equivalent to maximizing the speed.

From the steady<state solutionof previous example, the travelling time of the fastest solution is only 2.55% shorter than that of the ma.ximum constant speed solution, which has smooth control torque function. Therefore, if there exists such circlelike arc in the specified path, the maximum constant speed soluti<>n is a nice choice. In the example below, there only exists a short part in the travelling path with strict speed limit, a result from the trade-off between smoothness and travelling time is shown in f#4. Since s is chosen to be unit arc length, v(t) presents the travelling speed of Laser spot in the path space. Therefore, if Laser power is on-line adjustable, v(t) is the reference.5··specified path . $:--0· .. ·s = 23.1112 ... ;Travelling speed 5 r----~.....::..-"'----_,>-o>10xo5 time10_~Iqqcon~0510>time10s20~I qqcon$7Q d2 o.

$:--0· .. ·s = 23.1112 ... ;Travelling speed 5 r----~.....::..-"'----_,>-o>10xo5 time10_~Iqqcon~0510>time10s20~I qqcon$7Q d2 o. 5 time10fig#4 The trade-off between smoothness and travelling time: tf = 11.9221132If the travelling speed is limited above by the available Laser power, it is straightforward to put an extra constraint in phase plane and modify the control trajectory. This is shown in fig#5.~~ 5>time10s202r--------.---.......-+---, o -2L...--------------l 5 timeo10fig#5 Speed limited by available Laser power: tf = 12.3217If the comer is too sharp, that is the radius is too small, the speed. limit region in phase plane may be very low. Then, the Laser spot must slow down when pass that comer. In order to gain uniform degree ofsintering, Laser power need to be turned down. Thus, the speed is limited below by the continuity of sintering which does not occur if Laser power is too low.

In order to gain uniform degree ofsintering, Laser power need to be turned down. Thus, the speed is limited below by the continuity of sintering which does not occur if Laser power is too low. When, the radius and maximumtorque are given, the fastest speed around the comer could be seen in phase plane.•The design of the. radius. at comer should compromise with the speed requirement, available Laser pOwer. The change of infeasible speed region in phase plane corresponding to various radius is shown in fig#6.Specified path -~~.::-.::::: ~. -... ~ ·~ R"'I.4 '. . "',,', R= : R 6,· \..:... 5 .. .;·· .. ·.......... -: . : \ ' Ii Ii , . : I . . Or":" ··············i···.,. """R=8: rr': , , : ."-",~>oX510 s20fig#6 Infeasible speed region vs Radius of comer133ConclusionGiven a scanner system, the maximum torque, the specified path, and other circutnstance in SFFapplication,thequ~sti()n,'hQwJastcouldth~scannel"t>edriven 1', is answered in this paper.

It is nice to formulate the minimumitime optimalcontrol problem in s-space since (1) fhe problem is one-degree-of-freedom such that thephase~plane technique is useful, (2) v=ds/dt presentsthe travelling speed along the path which could be referred for any speed constraints when s is unit arc length. The control strategy is proposed. If the path is near circle .arc, the infeasible speed region is like that shown in fig#2 and fig#3, the maximum constant speed solution is chosen since it only needs 2.55% more than the time needed by optimal solution, and has smoother control trajectory. If speed is limited by available Laser power, the constraints could be straight forward added into phase plane diagram. If continuity and uniformity of sintering are required, the radius of comer in the path must be restricted.

If continuity and uniformity of sintering are required, the radius of comer in the path must be restricted. In phase plane diagram, it is easy to see the speed limit around the comer corresponding to the given radius.References[1][2][3][4][5][6][7]Bryson, A.E., Jr., and Ho, Y.C., 'Applied Optimal Control', Blaisdell Publishing Co., Waltham, MA, 1969. Bobrow, J.E., Dubowsky, S., and Gibson, J.S., 'Time-Optimal Control of Robotic Manipulators along Specified Paths', The international Journal of Robotics Research, Vo1.4, No.3, Fall 1985, p3-p17. Shin, KangG.,.and McKay, Neil D., 'Minimum-Time Control of Robotic Manipulators with Geometric Path Constraints', IEEE Transaction on Automatic Control, Vol.AC-30, No.6, June 1985, p531~p541. Stoer, 1., and Bulirsch, R., 'Introduction to Numerical Analysis', Springer-Verlag, New York, 1980.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCONTOURED 3D PRINTING OF FIBER REINFORCED POLYMERSL. Y. L. Tse*, S. Kapila†, and K. Barton*Department of Mechanical Engineering, University of Michigan, Ann Arbor, MI 48109 †Department of Orofacial Science, University of California, San Francisco, CA 94143AbstractAdditive manufacturing (AM) is a cost effective approach for small-scale production, providing higher design flexibility and less material waste than traditional manufacturing techniques. Polymers constitute one of the most popular AM materials, yielding lightweight but inherently weak components that cannot hold up against high tension and bending stresses. A need for improved tensile strength has driven a recent interest in AM of fiber reinforced polymers (FRPs).

A need for improved tensile strength has driven a recent interest in AM of fiber reinforced polymers (FRPs). AM-FRPs reinforced with short fibers have demonstrated increased mechanical strength, but with limited design and structural flexibility. AM-FRPs reinforced with continuous fibers provide structural reinforcement within plane; as such, the fibers cannot be extruded along a contoured profile, significantly minimizing the application space for these AM-FRP devices. In this article, we address this current gap through the development of a new FRP additive manufacturing process that is capable of continuous fiber deposition along contoured trajectories. Experimental demonstrations validate the proposed process.IntroductionAdditive Manufacturing (AM) is a relatively new manufacturing concept that is driving a new paradigm in advanced manufacturing [1-3].

By continuously attaching small quantities of material at precisely controlled locations, additive manufacturing allows the user to fabricate intricate geometries using polymer, metal, ceramic, or biological tissues and materials. Unlike many traditional manufacturing techniques, AM processes do not require templates, molds or masks; thus, AM leverages the distribution of material at specific locations to achieve unique device designs with time and cost savings for small to medium-scale production. Additionally, AM has the unique ability to fabricate interlocking geometries, features embedded within a shell structure, and heterogeneously printed materials within a single layer or design, leading to new designs and functionalities [4].Polymer inks were among the first materials utilized in AM for prototyping functional devices [4], and have remained the most commonly used material to date.

Recent advancements in material properties and AM processing capabilities have led to new AM materials for a wider range of applications. Flexible polymers such as polycaprolactone (PCL) and Nylon have recently been demonstrated for 3D printing [5,6]; these materials result in lightweight and geometrically customized parts, such as prosthetic limbs and robotic frames. Despite the advantages for customized prosthetics and robotic systems [7,8], the tensile strength of these materials is significantly lower than standard metal materials or that required for optimal functionality [9].Current advancements in polymer materials have resulted in fibers (specifically carbon fibers) being combined with a polymer into a single material to generate fiber-reinforced polymers (FRPs or CFRPs for carbon fiber reinforced polymers).

The use of fibers embedded within the1205polymer layers increases the strength-to-weight ratio and improves resistance against tension versus standard polymers (see https://markforged.com/materials/). Current FRP parts are fabricated through traditional approaches that involve stacking up layers of fibers and polymer on a mold or template, and applying heat or chemical curing to adhere the different layers together. This method has resulted in parts with significant increases in strength along the direction of the fibers. One common use of FRPs is towards the development of shell-structures for automotive applications. Despite the advantages in structural strength, this fabrication process is time consuming, expensive, and inflexible to design changes; a new mold is needed for each design change.To address these fabrication limitations with FRP, AM of FRPs has garnered recent significant interest.

Two common techniques for AM of FRP include: 1) printing of a mixture containing polymer and short fiber fragments using a standard Fused Deposition Modeling (FDM) technique [10], and 2) repeatedly printing alternative layers of polymer and continuous fiber with the FDM approach (Markforged printers). A core disadvantage of the first approach stems from the weak physical bonding between the short fibers. Short fibers do not produce significant reinforcements within the material as they lack a structured orientation and continuity against stress and strain forces. In contrast to short FRPs, continuous FRPs provide a large resistance force against tension and large bending strains, particularly along the optimum orientation due to the aligned positioning of the fibers.

Despite the potential advantages of this approach, continuous FRP printing has yet to be translated into a commercially viable additive manufacturing technique.Markforged has developed the first and currently only commercialized continuous FRP 3D printers. Structurally, Markforged printers are very close to that of regular FDM machines. The system prints a layer of polymer using a traditional FDM approach in which the polymer undergoes a phase transition (thermally melted) prior to being deposited onto the substrate surface. Following the polymer deposition, strands of continuous fiber are placed onto the melted polymer surface in rastered or looped patterns. A significant disadvantage of the current printer design is the restriction to 2D printing of fiber layers due to (1) the two-step layering process that requires at least one layer of polymer to be deposited prior to the fiber strands, and (2) bulky and rigid printheads that prohibit contour following.

As such, true 3D deposition of contoured fiber reinforced polymers has yet to be realized, prohibiting FRP shell-structures from being 3D printed.To address this important technical gap in 3D printing of FRPs, the authors have developed a new 3D printing methodology that enables 3D contoured deposition of FRP. This manuscript presents a 3D printing system that enables the placement of continuous FRP filaments onto contoured surfaces (e.g. XYZ fiber placement in real-time), thereby allowing the user to create shell-structures similar to the structures fabricated with traditional manufacturing approaches. One important advantage of 3D printing these structures is the ability to customize the designs on a case-by-case basis.The manuscript is arranged as follows. Section II provides some background information on FDM and the application needs for additive manufacturing of FRPs.

Section II provides some background information on FDM and the application needs for additive manufacturing of FRPs. Section III introduces the design and development of a critical element in the printing process known as the co-axial fiber, which is a polymer-coated fiber that provides rigidity and adherence within the combined structure. Section IV provides additional details of the printhead including the unique features in the design that have led to the development of true 3D printing of FRPs. A demonstration of the1206capability of the printhead is presented in section V. A discussion of design challenges, printing process limitations, and future development directions is given in section VI. Section VII contains concluding remarks.BackgroundA. Fused Deposition Modeling (FDM)Conventional FDM follows the additive paradigm of depositing material in successive layers in order to produce a 3-Dimensional (3D) part.

Fused Deposition Modeling (FDM)Conventional FDM follows the additive paradigm of depositing material in successive layers in order to produce a 3-Dimensional (3D) part. Contrary to powder or ink-based deposition processes, FDM extrudes flattened lines of molten material to form layers that rapidly solidify upon extrusion from the nozzle. Commonly available FDM materials, such as Acrylonitrile- Butadiene-Styrene (ABS) and polylactic acid (PLA), are thermoplastics; that is they become molten when heat is applied to them and cool rapidly in air after the heating source is removed. Moreover, their molten forms are viscoelastic fluids that enable these materials to sustain their own weight and shape upon release from the nozzle, resulting in printed filaments with ~1:1 aspect (height-to-width) ratio.

By overlapping 2D patterns printed with filaments of ~1:1 aspect ratio, the FDM process can create very complicated 3D structures that are challenging to produce with traditional manufacturing techniques such as lithography or micromachining.As described previously, existing FRP printers apply a conventional FDM approach with either continuous or short fibers. The short FRPs are not as flexible and lack an optimum loading orientation, while continuous FRP can only align fibers in planar directions. Importantly, the ability to deposit contoured shapes cannot currently be achieved using the conventional FDM printers.B. Applications for FRPFRPs that have optimal shape and properties have a multitude of potential applications in the automotive, aerospace and medical fields.

Applications for FRPFRPs that have optimal shape and properties have a multitude of potential applications in the automotive, aerospace and medical fields. To meet the rising regulations for environment protection [11] and safety standards [12, 13], there is a growing demand for FRPs that reduce the weight of automobiles, while simultaneously making their structures more effective in shock energy absorption [14]. Additionally, as more alternative fuels are being adopted in the automotive industry [16], there is an increasing need for materials that withstand the unique conditions created by these fuels within the power-train.

These developments have led to an increasing demand for new manufacturing technologies that enable the fabrication of FRP-based power-trains, battery casings and thermal management systems that are capable of withstanding the harsh environments introduced by these different fuels.In the medical field, assistive devices, in particular ankle-foot orthoses (AFOs), artificial limbs and increasingly robotics, are custom-made medical devices that correct for a variety of acquired or congenital defects in patients. The current, widely-adopted, manufacturing approach for fabricating AFOs is a plaster casting method that requires multiple fittings, a costly and time- intensive fabrication process, and results in repeat visits for adjustments and additional modifications [17].

Additive manufacturing of FRPs provides an alternative fabrication approach that has the potential to decrease production times, lower waste and energy consumption, decrease costs, and improve AFO customization and user performance [18-20].1207In the following sections, we will discuss the materials, features and design challenges of our FRP printing system.Co-axial FiberAn important difference between existing FRP 3D printers and the system described herein resides in the use of a spool of polymer coated fiber filament known as the co-axial fiber, as opposed to adhering fiber threads onto a polymer foundation used in currently available FRP printers. The spool of material is fed through a nozzle, while simultaneously melting the outer coating and adhering the filament fibers to previously printed material.

The spool of material is fed through a nozzle, while simultaneously melting the outer coating and adhering the filament fibers to previously printed material. This approach provides three important advantages over existing technology and methods: (1) the more rigid structure of the co-axial fiber enables the printer to deposit material along contoured surfaces and achieve overhanging features, which have not yet been realized with conventional FRP printing; (2) the one-step approach provides faster fabrication with less material waste; (3) the co-axial fiber provides important structural advantages such as the reduction of buckling during fabrication as compared to fiber threads currently in use.A. Raw MaterialsThe raw materials in this process consist of the fiber and polymer materials that are used to make-up the FRP strands.

Raw MaterialsThe raw materials in this process consist of the fiber and polymer materials that are used to make-up the FRP strands. Printing demonstrations in this manuscript utilize carbon fibers due to the ubiquitous use of this material in practice; however, potential fiber materials can be expanded to include Kevlar, glass fibers, metal wires, and medical suture materials.The polymer base used in the manuscript consists of Polycaprolactone (PCL). Additional polymer options may include nylon and ABS, among others. As discussed previously, FDM compatible polymers exhibit a viscoelastic molten form and rapid solidification property; thus allowing the polymer to maintain a ~1:1 aspect ratio during deposition. The approach presented here utilizes co-axial fibers that maintain the 1:1 aspect ratio through the fiber thread rather the polymer.

The approach presented here utilizes co-axial fibers that maintain the 1:1 aspect ratio through the fiber thread rather the polymer. Therefore, candidate polymers do not have to exhibit a viscoelastic molten form or rapid solidification this property; polymer compatibility with printing system is significantly more diverse than conventional FDM systems.thanB. Co-axial Fiber FabricationThe co-axial filament is fabricated using the set-up depicted in fig. 1. The setup consists of a heat sink, a heating element, a temperature controller, a syringe barrel that holds a reservoir of molten polymer (PCL), a frame structure, a motor and a spool core to collect the co-axial fiber filament.Fig. 1. Schematic of filament fabricationThe coating process requires the carbon fiber to be continuously pulled through the molten1208PCL in the syringe barrel using a motor that drives the spool core.

The co-axial fiber is then released from the bottom of the syringe and cooled down during the transport to the spool core, where it is wound for use in the printing process. The diameter of the filament depends on: (1) the diameter of the fiber threads; (2) the inner diameter of the extrusion needle; and (3) the speed at which the filament is being pulled. For each needle diameter, there is a range of optimum motor rotational speeds that ensure consistent polymer coating; high speeds lead to thin and uneven polymer coatings, while slow speeds result in the formation of polymer beads along the filament.For the co-axial fiber fabrication setup in fig. 1, the diameter of the fiber threads is ~0.4mm, the extrusion needle inner diameter is 1mm, and the filament is being pulled at a speed of ~0.9mm/s. The resultant co-axial fiber has a diameter of ~ 0.9mm.

The resultant co-axial fiber has a diameter of ~ 0.9mm. To improve the printed device resolution, the co-axial fiber diameter can be modified by reducing the number of fiber threads within the filament, or decreasing the inner diameter of the extrusion needle. Co-axial fiber diameters that are feasible using the materials demonstrated in this paper range from 0.4mm to 1.2mm.Printhead DesignA major consideration in our printhead design focuses on the development of a long, slender profile to enable 3D printing along highly articulating profiles. The design consists of three main sub-systems: (1) an actuation system, (2) an anchoring system, and (3) an extraction system. Each sub-system is driven with a stepper motor. Unlike common FDM printhead designs that incorporate a motor at the printing tip, a slender profile requires additional mechanisms to transfer the motor action to the printing tip.A.

Unlike common FDM printhead designs that incorporate a motor at the printing tip, a slender profile requires additional mechanisms to transfer the motor action to the printing tip.A. Actuation systemThe actuation system consists of a driving motor with gripping wheels, a long metal tube that directs the co-axial fiber to the extraction system, and a nozzle that deposits the co-axial fiber filament onto the printing surface (see fig. 2). The driving motor is a NEMA 17 motor with a convention FDM gripper wheel. A gripper wheel with a rubber outer lining provides a flexible surface for effectively directing the co-axial fiber.Fig. 2. Actuation system: Gripper wheels push the co-axial fiber into the aligment tube and out through the nozzle.The two gripping wheels rotate as the motor turns, sending the polymer coated filament into the long metal tube. The metal tube is ~150mm long with an inner diameter of 2mm.

The metal tube is ~150mm long with an inner diameter of 2mm. The tube deposits the co-axial fiber directly below the extractor (cutting blade) that is elevated until actuated. The polymer coating of our co-axial fiber enables the fiber to traverse the 150mm long tube without becoming tangled or buckling, which is a common occurrence with uncoated carbon fiber bundles.The co-axial fiber enters a bent nozzle after exiting the alignment tube. The role of this nozzle is to bend the filament with a ~5mm radius, therefore allowing the filament to be deposited tangentially to the substrate rather than in a perpendicular orientation. This reduces clogging and1209misalignment issues that occur due to the sudden change in direction from perpendicular to tangential when the co-axial fiber is deposited using a vertical orientation.B. Filament anchoring systemThe filament anchoring system is comprised of a motor (NEMA 17), crank shaft, and heating pad (see fig. 3).

Filament anchoring systemThe filament anchoring system is comprised of a motor (NEMA 17), crank shaft, and heating pad (see fig. 3). As the motor turns, the crank shaft (pivoted in the middle) will transfer the rotational motion of the motor shaft to the heating pad. The heating pad is heated with loops of copper wires charged with a 5V DC supply, and can move in both clockwise and counter- clockwise directions. The heating pad supplies a small thermal source to heat the polymer coating of the co-axial fiber, and enable the fiber to adhere to the substrate surface. The rotational movement from the crankshaft combined with the compression from the springs and flexible heating pad allows the printhead to deform elastically as required by the substrate geometry.Fig. 3.

3. Filament anchoring system: A cyclical motion aids in compression and adherence of the co-axial fiber onto the substrate.A major challenge in the filament anchoring system stems from the adhesion challenges with the co-axial fiber filament. When the polymer coating of the co-axial filament melts, the molten polymer coating becomes an adhesive agent. As the temperature of the filament decreases, the adhesive nature of the molten polymer reduces. As one might expect, the molten polymer has a stronger adhesion to the higher temperature heating element as compared to the substrate. To minimize the adhesion forces between the polymer and the heating pad, a rocking motion has been introduced with the crankshaft, enabling the heating pad to transfer the co-axial fiber to the substrate.It is important to note that the heating element in our design is not attached to the nozzle, a common design approach in many FDM printing systems.

This separation between the heating element and the nozzle is critical to ensure that the polymer coating of the co-axial fiber filament does not melt within the nozzle and cause the system to clog.C. Extraction systemThe extraction system contains a NEMA 17 motor with a 5:1 gear-box, a rotational-to-linear motion conversion mechanism and a pivoted cutting blade (fig. 4). The gear- box provides a torque multiplier to drive the cutting blade with a strong force. The rotational-to-linear motion conversion mechanism also provides a multiplier to the force generated by the motor. The cutting blade is modeled after a nail cutter (fig. 4) with a pivot screw localizing cutting force along the blade.Fig. 4. Extraction system: A lead screw converts rotational motion into linear motion to provide the cutting force.To initiate the extraction process, the motor drives the cutting blade onto the co-axial fiber,1210severing the filament along the blade.

Once the co-axial filament has been severed, the driving motor from the actuation system recoils the co-axial cable to ensure a clear separation between the two severed ends. In the last step, the cutting blade is retracted into a raised position while the co- axial fiber returns to the tip of the nozzle, thereby resetting the extraction system.Importantly, the extraction system must be activated approximately 10mm prior to the desired location. The delay is a function of the separation between the extraction location and the printing tip. The remaining co-axial fiber is passively deposited by the anchored filament onto the printing surface.D. Printer setupFigure 5 provides an image of the FRP system located in the authors’ lab. The 4-degrees of freedom (4-DOF) system consists of Aerotech X, Y, Z and rotary motor stages with an XY resolution of 1 micron, Z resolution of 2µm, and rotatory stage with a 1-degree resolution.

The 4-degrees of freedom (4-DOF) system consists of Aerotech X, Y, Z and rotary motor stages with an XY resolution of 1 micron, Z resolution of 2µm, and rotatory stage with a 1-degree resolution. The printhead is mounted to the Z-axis stage, while the substrate is mounted to the rotatory stage mounted on the XY stage. The printhead has been designed to handle tilted surfaces in the YZ plane up to 65o in the positive and negative direction from the horizontal. Although the current design can only follow contours in the YZ plane, future designs will enable arbitrary contour following. The combined motion of the 4 stages enables the printhead to deposit complicated 3D patterns of co-axial fibers on the printing surface.Fig. 5.

The combined motion of the 4 stages enables the printhead to deposit complicated 3D patterns of co-axial fibers on the printing surface.Fig. 5. Custom FRP printer with 4-DOFPrinting DemonstrationsTo demonstrate the complex deposition capabilities of this FRP printer and highlightgeometries that cannot be printed with current fiber reinforced polymer printers, this section presents three parts printed on support Polyvinyl Alcohol substrates. PVA was selected as the substrate / support structure material due to its ability to be dissolved in water, and thus separated easily from the printed FRP structure.(PVA)Fig. 6. CAD drawing of the bi-planar PVA substrate and desired co-axial fliament printing path. Curved loops (red lines) were followed by repetitive rasters (black lines).A. Bi-planar printingA CAD drawing of a bi-planar (YZ, XY) PVA support substrate and co-axial fiber deposition path is provided in fig. 6. The outer, bi-planar lines (red lines in fig.

Bi-planar printingA CAD drawing of a bi-planar (YZ, XY) PVA support substrate and co-axial fiber deposition path is provided in fig. 6. The outer, bi-planar lines (red lines in fig. 6) are printed first followed by the more conventional, rastered lines along the contoured (YZ) surface (black lines in fig. 6). While either of the planes could be oriented to enable conventional FRP printers to fabricate the lines, the combined pattern1211Fig. 7. a). Printing of a bi-planar FRP part on a PVA substrate. b). Completed part top view. c). Completed part bottom view. d). Pre-bent printed FRP part. e). Printed FRP part under bending stress. Part elastically deforms under bending stress and springs back to the original shape when the stress is removed.contains seamless transitions between the planes, thereby requiring printing along a contoured surface. This pattern represents the first demonstration of contoured, 3D FRP printing.

This pattern represents the first demonstration of contoured, 3D FRP printing. Figure 7a shows the FRP printing in progress, while figs. 7b and 7c show the completed part after it has been removed from the PVA substrate base. Single layer co-axial fiber patterns (0.9 mm diameter lines, 1 mm layer thickness) were found to be flexible and easily extracted from the PVA substrate base. Importantly, the mechanical behavior demonstrated in fig. 7e fabricated using follows conventional manufacturing methods as compared to FRP components fabricated on FDM-style printers.that of FRP componentsFig. 8. CAD image of a multi- planar PVA substrate with co- axial fliament printing path.B. Multi-planar printingFigure 8 shows a CAD drawing of a multi-planar PVA support substrate and the path of the co-axial fiber. Eight loops of co-axial fiber were printed on the multi- planar support substrate. Figure 9a shows the FRP printing in progress, while fig.

Eight loops of co-axial fiber were printed on the multi- planar support substrate. Figure 9a shows the FRP printing in progress, while fig. 9b shows the completed FRP pattern after it has been removed from the multi-planar PVA substrate base.The design in fig. 8 consists of repeating curves in the YZ, XZ and XY planes. This particular architecture was chosen as it represents a simple design that cannot currently be fabricated using existing FRP printing systems. Similar to the elastic deformation demonstrated with the bi-planar part, fig. 9c illustrates the deformation capabilities of the multi-planar part.Fig. 9. a). Multi-planar printing of FRP part on a PVA substrate. b). Multi-planar part unstretched. c). Stretched multi-planar part.1212Fig. 10. a). Multi-layered, multi-planar FRP printed pattern on PVA substrate. b). Multi-layered, multi-planar printed part after being removed from the PVA substrate.

10. a). Multi-layered, multi-planar FRP printed pattern on PVA substrate. b). Multi-layered, multi-planar printed part after being removed from the PVA substrate. The five layers enhance strutural strength and decrease elastic deformation.C. Multi-layered, multi-planar printingUtilizing the substrate architecture from fig. 8, this section demonstrates a multi-layered, multi-planar printed 3D-FRP pattern. Each of the five layers consists of 8 curved loops (~60 mm diameter), with the layers deposited in an overlapping fashion. To ensure adhesion between the layers, the co-axial fibers were pressed into the previously deposited layer. This resulted in layers that were slightly smaller than the single layered pattern (~2.7 mm across 5 layers). The printed part on the PVA supporting substrate is shown in fig. 10a.

This resulted in layers that were slightly smaller than the single layered pattern (~2.7 mm across 5 layers). The printed part on the PVA supporting substrate is shown in fig. 10a. The addition of multiple layers resulted in a much stronger adhesive force between the substrate and the printed pattern, requiring the substrate to be dissolved in water to retrieve the printed FRP pattern (fig. 10b). Additionally, the extra layers significantly decreased the elastic deformation characteristics of the printed pattern.DiscussionThe experimental results presented in figs. 7, 9 and 10 provide the first demonstration of contoured, multi-layered additively manufactured FRP parts.

7, 9 and 10 provide the first demonstration of contoured, multi-layered additively manufactured FRP parts. While the fabrication of contoured, multi-layered FRP parts has been demonstrated using conventional methods such as stacking layers of FRP sheets within a mold, the ability to achieve real-time design customization through additive manufacturing presents a unique capability that will expand the application of FRP in commercial technologies.One of the most important innovations of our FRP printing system stems from the use of a co-axial fiber that enables fiber deposition in 3-dimensions (XYZ planes) within the same build. The co-axial cables provides the structural support necessary to guide and adhere fibers out of the XY plane; existing FRP 3D printers can only deposit fibers in the XY plane. This additional functionality enables contoured and 3-dimensional FRP printing for the first time.

This additional functionality enables contoured and 3-dimensional FRP printing for the first time. The impact of this advancement will be two-fold: (1) Customized 3D printing of contoured FRP components will enable the use of FRPs in applications ranging from the automotive and aerospace industries to the biomedical field. (2) The co-axial fiber provides a unique structural support mechanism that can be leveraged during the fabrication process to customize the strength-to-weight ratio, and provide previously unattainable 3D printed design features such as overhanging features, contoured shell structures, and 3-dimensional shapes.Despite the success of this initial prototype, there are important design modifications that will further enhance the fabrication capabilities of our FRP printing system. For example, a more compact printhead design across all subsystems will enable the fabrication of higher resolution patterns within a broader design space (e.g.

For example, a more compact printhead design across all subsystems will enable the fabrication of higher resolution patterns within a broader design space (e.g. right-angles, 360 degree continuous depositions).1213Additionally, to extend 3D printing of contoured FRP patters to commercial applications will require a faster, more flexible, cost effective, and robust printing system. Future FRP printing system designs aim to address this need.ConclusionThe use of fiber reinforced polymer materials has been growing within the automotive, aerospace and biomedical industries, driving the development of corresponding 3D printing technologies. Currently FRP printing techniques are unable to align fibers along a contoured surface, therefore preventing them from creating shell structures with similar strength-to-weight ratios as those made with traditional layering/mold approaches.

The FRP printing system detailed in this manuscript provides the first demonstration of 3D printed FRP patterns with the fibers aligned along the contour profile. This represents a significant step towards customized 3D printing of functional contoured structures. Additionally, advancements in the printhead design will address the need for a faster, more flexible and robust printing system. Finally, we will scale up the printing system to extend the applications to the automotive and aerospace industries.References[1] K. V. Wong, and A. Hernandez, “A review of additive manufacturing,” ISRN MechanicalEngineering, 2012.[2] S. H. Huang, P. Liu, A. Mokasdar, and L. Hou, “Additive manufacturing and its societalimpact: a literature review,” Int J Adv Manuf Tech, 67(5-8), pp. 1191-1203, Jul. 2013.[3] W. E. Frazier, “Metal additive manufacturing: a review,” J Mater Eng Perform, 23(6), pp.1917-1928, Jul. 2014.[4] I. Gibson, D. Rosen and B.

ntrol Methods for the Electron Beam Free Form Fabrication ProcessW. J. Seufzer, K. M. Taminger NASA Langley Research Center, Hampton VA 23681AbstractEngineering a closed-loop control system for an electron beam welder for additivemanufacturing is challenging. For earth and space based applications, components must work in a vacuum and optical components must be protected from becoming occluded with metal vapor. For extraterrestrial applications added components increase launch weight and increase complexity. Here we present three different control methods for electron beam free form fabrication. A relatively simple coarse feedback control method is introduced that couples path planning and electron beam parameter controls into the build process to increase flexibility and improve build quality.

The different approaches may be applied separately or together to provide enhanced EBF3 system performance.IntroductionResearchers at NASA Langley Research Center have developed the Electron Beam Free Form Fabrication (EBF3) process, a rapid metal deposition process that works efficiently with a variety of weldable alloys [1,2,3]. The EBF3 process can be used to build a complex, unitized part in a layer-additive fashion, although the more immediate payoff is for use as a manufacturing process for adding details to components fabricated from simplified castings and forgings or plate products. Figure 1 shows a schematic of an EBF3 system. The EBF3 process introduces metal wire feedstock into a molten pool that is created and sustained using a focused electron beam in a vacuum environment (1x10-4 torr or lower).

The EBF3 process introduces metal wire feedstock into a molten pool that is created and sustained using a focused electron beam in a vacuum environment (1x10-4 torr or lower). Operation in a vacuum ensures a clean process environment and eliminates the need for a consumable shield gas.Vacuum chamberElectron beam gunWire feedElectron beamDepositzSubstratePositioning systemxyFigure 1. Schematic of electron beam freeform fabrication (EBF3) system.The EBF3 process is nearly 100% efficient in feedstock consumption and approaches 95%efficiency in power usage. The electron beam couples effectively with any electrically conductive material, including highly reflective alloys such as aluminum and copper. A variety of weldable alloys can be processed using EBF3; further development is required to determine if non-weldable alloys can also be deposited.

A variety of weldable alloys can be processed using EBF3; further development is required to determine if non-weldable alloys can also be deposited. The EBF3 process is capable of bulk metal deposition at deposition rates in excess of 150 in3 hr-1 as well as finer detail at lower deposition rates with the same equipment. The diameter of the wire feedstock is the controlling factor determining the smallest detail attainable using this process: fine diameter wires may be used for adding fine details and large diameter wires to increase deposition rate during bulk deposition.

In a system with dual wire feed that can be controlled simultaneously and independently, the two wire feeders may be loaded with either a fine and a coarse wire diameter for different feature definition, or two different alloys to facilitate producing components with compositional gradients.As with all new solid freeform fabrication processes, much of the control and selection of processing parameters for the EBF3 process have been empirically derived. Better understanding of the process is required to enable development of an automated control system. Lessons learned and techniques developed for another metal deposition process, the Laser Engineered Net Shaping (LENSTM) process, may be applicable to the EBF3 process [5,6]. The LENSTM closed loop feedback control was granted US Patent number 6,459,951, October 1, 2002, and describes a method using optical and thermal imaging to monitor the molten pool and deposition height to achieve a closed loop process.

However, due to difficulties with controlling the mass flow rate of the powder and slow update rates of the lamp-pumped Nd:YAG laser in the LENSTM process, control was established indirectly through translation speed control governed by the understanding of the effects of input process parameters on the thermodynamics (as indicated by the melt pool size and shape) and geometry of the deposit (as indicated by the bead height). The EBF3 process offers many additional degrees of freedom in directly controllable variables as compared to that of the LENSTM process, which permits development of a different closed loop control methodology and ability to achieve finer process control.Control MethodsInput parameters for the EBF3 process include beam power, beam pattern, travel speed, and wire feed rate. Variations in these parameters influence the deposit height and width (geometry), alloy composition (chemistry), residual stresses within the final part, and distortions of the final part.

Variations in these parameters influence the deposit height and width (geometry), alloy composition (chemistry), residual stresses within the final part, and distortions of the final part. The input parameters have been shown to have coupled interactions that currently require trial and error, and expert human experience, to find desirable combinations. Even with a set of parameters that provide an acceptable deposit geometry, chemistry, etc., there is still some variability within the process to motivate the design of a closed-loop control system. This system would be able to monitor the build process, measure the variability, and correct the system as the build progresses.There are three methods to consider for EBF3 process control: feed forward, coarse feedbackand fine feedback. Each of the three process controls is explained here along with theirrespective advantages and disadvantages.

Each of the three process controls is explained here along with theirrespective advantages and disadvantages. Since EBF3 can be applied to industrial, low Earth orbit, and Lunar/Mars surface operations, design challenges for those applications are also discussed.Feed ForwardThe term "feed forward" is used to indicate that information about the deposition process is well known a priori. That is, the information only flows forward during the build process, there is no feedback. Figure 2 illustrates the steps used in the process. First, CAD (Computer Aided Design) software is used to define the part geometry. Second, CAM (Computer Aided Manufacturing) software is used to reduce the part definition into commands to build layers in the additive process. Third, the EBEAM step is used insert electron beam control parameters into the CAM file.

Third, the EBEAM step is used insert electron beam control parameters into the CAM file. Fourth, CNC (Computer Numeric Control), electron beam control parameters and tool path commands, are given to the electron beam welder to produce the part.Figure 2. The classic approach used for subtractive manufacturing with an additional step, EBEAM, needed for EBF3 additive manufacturing. The arrows represent a step-by-step process with information flow only in the forward direction.The non-linear relationships between input parameters and their role in a deposits final chemistry was shown in a recent DOE (Design Of Experiments)[4]. Figure 3 shows travel speed and beam power each taken individually play a very small role in deposit chemistry. Collectively they play the second strongest role in characterizing the deposit chemistry indicating a strong coupling between these 2 input parameters on that particular output parameter.Figure 3.

Collectively they play the second strongest role in characterizing the deposit chemistry indicating a strong coupling between these 2 input parameters on that particular output parameter.Figure 3. Design of Experiments results showing the complex interaction of wire speed, beam power, and travel speed and their impact on Aluminum chemistry in Ti-6Al-4V.The main advantage of a feed forward process control is its apparent simplicity since the electron beam welder, as delivered from its manufacturer, needs no modification. A disadvantage to this method is that a human expert trained through trial and error must select the input parameter combination. Additionally, even with a good selection of electron beam control parameters, only a few simple parts have been built without operator intervention. That is, a human in the loop watches the build process and makes small adjustments to deposition parameters as the build progresses.

That is, a human in the loop watches the build process and makes small adjustments to deposition parameters as the build progresses. Further, even with operator adjustments, errors may arise. Figure 4 shows a build with an errant corner. About 50 layers were deposited very consistently but near the end of the build, something changed and the small error on one layer became larger in subsequent layers. In other builds attempts have been made to correct problem builds with human intervention with varying levels of success.Figure 4. A build error after about 50 successful layers. A small error in one layer becomes larger in subsequent layers.The feed forward method can be improved through experiments and DOE analysis. Data from those experiments would be used to quantify the relationships among the parameters and make it possible to capture that process knowledge into algorithms.

Data from those experiments would be used to quantify the relationships among the parameters and make it possible to capture that process knowledge into algorithms. These relationships need to be understood for the feed forward method, and for any other type of control method. Open loop deposition with human process control has been used to build a variety of different shapes, but repeatability and quality control is extremely difficult to achieve. The feed forward approach helps to improve the repeatability and build quality, but is unable to identify and correct errors, prevent propagation of errors, and ensure part accuracy.Coarse feedbackCoarse feedback control is an intermediate step that closes the control loop, not at the level ofreal-time control, but a broader level. The steps from the classic CAD/CAM/CNC process are still used but a feedback path and a comparison function are added. In this method, the original part drawing is included in the control loop.

The steps from the classic CAD/CAM/CNC process are still used but a feedback path and a comparison function are added. In this method, the original part drawing is included in the control loop. Figure 5 shows the information flow. First, one layer is taken from the part definition (CAD) and used to generate a tool path (CAM). EBF3 process parameter information is computed and added to the tool path and the welder (CNC) builds that layer. After one, or several layers (depending on tolerance requirements), the height profile of the deposit is taken. That height information is fed back into the system and is used to compute what layer height will be taken next from the CAD description. For instance, if a deposit bead height of 0.030 in. was assumed (based on feed forward experiments) but an average beadheight of 0.025 in. was measured, the next layer from the CAD drawing will be taken from 0.025 in. above the previous layer. This 0.005 in.

was measured, the next layer from the CAD drawing will be taken from 0.025 in. above the previous layer. This 0.005 in. error, if propagated over 100 layers, results in a part that is 0.5 in. too small, resulting in a 20% scaling factor only in the Z (height) direction.Figure 5. Adding a coarse feedback path to the feed forward approach. Here the direction arrows show a feedback path of information that will be used to plan the next layer of the build.Corrections are not limited to layer height. If an error occurs and a depression is measured along the tool path, the coarse feedback loop can be used to modify the EBF3 parameters for the next several layers until the depression is filled in. Additionally, corrections are not limited to height.

Additionally, corrections are not limited to height. Depending on the features of the post-build measuring system, if bead width can also be measured, corrections for bead width could also be applied.Any corrections that must be applied still rely on a thorough understanding of how the EBF3 parameters work in the feed forward method. For instance, a correction to bead height may not be a simple matter of tweaking one parameter over the measured distance. It appears necessary to tweak several parameters concurrently to affect the correct build geometry and chemistry. As such, it is not likely that a simple linear controller will control the EBF3 process.The main advantage of coarse feedback control is that it is a relatively simple developmentalstep from the feed forward method. Sensors and processing are needed to measure the build geometry, but these can be simple off-the-shelf measuring devices.

Sensors and processing are needed to measure the build geometry, but these can be simple off-the-shelf measuring devices. Once geometry information is obtained, comparing that information with the CAD diagram is a straight forward process. Any equipment used to monitor the build must work in a vacuum and survive the metal vapor atmosphere. Since measuring can be accomplished while the electron beam is off, shutters can protect optical devices from metal vapor deposition.The coarse feedback system offers several benefits for space based applications. First, additional equipment adds weight and directly influences launch costs; this method adds very little equipment. Second, even before launch, a simpler system is easier to integrate and test.

First, additional equipment adds weight and directly influences launch costs; this method adds very little equipment. Second, even before launch, a simpler system is easier to integrate and test. Third, a simple system should require less time for maintenance, lower energy requirements, and be more reliable.Fine feedbackFine feedback incorporates sensors to monitor and correct the deposition process in real-time.Figure 6 shows the flow of information for this method. As with the feed forward and coarse feedback methods, the first steps are to take the CAD drawing, slice it into layers with CAM software, and merge the tool path information with EBF3 control parameters. This method assumes that the feed forward information will yield a build that is relatively close to the desired geometry and chemistry and that any corrections will be relatively small.

This method assumes that the feed forward information will yield a build that is relatively close to the desired geometry and chemistry and that any corrections will be relatively small. As the electron beam welder is depositing a layer, the process is closely monitored and information is immediately available for feedback. The measured process features are compared with the desired build geometry and an error signal is used to tweak the CNC controller.Figure 6. Adding a feedback path to the classic approach. Here the direction arrows show a feedback path of information that will be used to correct the CNC controls in real-time. In comparison to the coarse feedback control the CAM information, or the layers of the build, are all computed once.Melt pool temperature, melt pool geometry, and the height of the build are the likelyparameters for monitoring the EBF3 process.

Direct observation of the melt pool reveals that as layers are added, residual heat from previous layers affects the melt pool geometry. Feedback to maintain a relatively constant temperature in the melt pool may affect the repeatability of the EBF3 process from layer to layer (for geometry and chemistry).Temperature and melt pool geometry are interrelated and monitoring temperature alone maybe sufficient to obtain consistent metallurgy. Measuring the geometry of the melt pool has implications for how the wire is approaching the melt pool. As wire approaches the melt pool it begins to melt in the electron beam and then flows into the melt pool. The surface tension of the liquid metal allows a bridge to exist between the approaching wire and the melt pool. Monitoring the shape of this bridge is an indicator of whether or not wire is approaching the melt pool at a desirable height.

Monitoring the shape of this bridge is an indicator of whether or not wire is approaching the melt pool at a desirable height. Observation in Earth-g, Lunar-g, Martian-g, and 0-g, reveal that surface tension, and therefore maintaining this bridge, is critical to the build process. Image processing could beimplemented to measure this bridge and provide a feedback signal to the circuit that controls the welder’s Z height.If melt pool temperature, melt pool geometry, and build height can be effectively measured and fed into the electron beam welder control system, another thing to consider is the response time of the electron beam welder components. The reaction times of the position motion axes, wire feed mechanism, and electron beam power, need to be assessed so that the control system design is stable.The advantage of the fine feedback method is that feedback about the build process is almostimmediate and should provide a very consistent build.

For example, there should be almost no height variability at a given layer such that the next layer would need to be modified. Also, with close monitoring and control of the melt pool, the chemistry of the build should have a narrower range of quality. The additional monitoring and computing equipment adds complexity to the system, but the complexity may be worthwhile if it enables unattended operation, enables repeatable process control, and increases build quality. While the additional complexity may justify itself on the industrial shop floor, it has disadvantages that are more severe for space- based applications. As mentioned earlier, complexity adds to system weight, increases integration and test requirements, potentially increases maintenance requirements, and potentially reduces reliability. The weight of the associated measuring and computational systems required for achieving coarse vs.

The weight of the associated measuring and computational systems required for achieving coarse vs. fine feedback control must be carefully examined to determine which approach provides the best trade-off between weight, complexity, and performance.SummaryThere are reasons for pursuing all three feed forward and feedback approaches. Serially theyrepresent a development pathway from a simple method to more complex methods. First, understanding the complex relationships between the input parameters and building that understanding into algorithms broadens the application to more than just a limited number of successful parameter settings. This has implications not only for complex build geometries but also chemistry, residual stress, and distortion control. Second, adding a coarse loop control to the system appears to be a simple developmental step that can be accomplished with off-the-shelf hardware and relatively simple software.

Second, adding a coarse loop control to the system appears to be a simple developmental step that can be accomplished with off-the-shelf hardware and relatively simple software. This developmental step also adds consideration for repair applications where the topographic features of part are measured and compared the original CAD drawing so that tool paths and EBF3 parameters are issued to add missing material. Selection of successful EBF3 parameters is dependent on work done for open-loop control. Third, adding fine loop control offers the possibility of higher quality control for geometry and chemistry, but with the cost of higher complexity. For some applications it may be beneficial to incorporate all three methods in unison, while others may derive adequate performance thru application of one or more methods discussed.References[1]Taminger, K. M. B.; and Hafley, R.

June 27, 2007.[5]Griffith, M. L.; et al.: “Understanding Thermal Behavior in LENS Processing of Structural Materials.” Proceedings of the TMS Annual Meeting, 1999.[6]Griffith, M. L.; et al.: “Understanding the Microstructure and Properties of Components Fabricated by Laser Engineered Net Shaping (LENS).” Department of Energy report. June 2000..

NTROL OF LASER CLADDING FOR RAPID PROTOTYPING – A REVIEWMallikharjuna R. Boddu, Robert G. Landers and Frank W. LiouDepartment of Mechanical and Aerospace Engineering and Engineering Mechanics, University of Missouri – Rolla, MO 65409AbstractLasers have wide–ranging applications in the manufacturing field (e.g., cladding, welding, cutting, machining, drilling). Extensive work is being conducted to apply laser cladding as a Rapid Prototyping (RP) process. In this paper the authors illustrate various principles of laser cladding in rapid prototyping. Important process parameters for the control of the laser cladding process are discussed as well as the experimental methods adopted, and results obtained by, various authors.IntroductionThe past decade has witnessed the emergence of new manufacturing technologies, where manufacturing time for parts of virtually any complexity is measured in hours, instead of days, weeks or months.

This is when Rapid Prototyping (RP) was conceived. Many RP technologies are available in the marketplace; however, these technologies utilize plastic (or similar) material to create parts and, thus, many parts are non–functional. Sterolithography Apparatus was the first commercially available RP system that successfully produced physical prototypes. It enabled the visualization of components produced directly from a CAD model by polymer curing with lasers. Laser–based RP systems have been introduced as a means of creating functional, metal prototypes with near–net shape geometries and development efforts are being conducted in research centers throughout the world (Laeng et al., 2000). The drawbacks to these systems are their low productivity and inability to consistently regulate part quality in terms of mechanical properties and geometry. To overcome these drawbacks, process control strategies have been utilized.

To overcome these drawbacks, process control strategies have been utilized. This paper provides an overview of this body of research.Lasers have provided industry with a new form of energy. Their wide application is due to their ability to act as a medium for communication, photography, and medical applications, as well as their ability to evaporate materials at the atomic level. Laser applications in manufacturing industries include welding, cutting, surface treatment, and ablation, and have recently extended their potential to RP. Laser-aided RP is a significant advance in traditional RP techniques due to the direct fabrication of a near-net shape part compared to the two step process involving an intermediate step of mould preparation in conventional RP techniques. Laser aided RP is advancing the state-of-the-art in product design by extending the laser cladding concept to RP.

Laser aided RP is advancing the state-of-the-art in product design by extending the laser cladding concept to RP. Some of the earlier attempts to build complex parts by layer addition were based on laser cladding principles. Studies have been carried out to determine the effect of process parameters on the quality of the clad layer (Weerasinghe and Steen, 1983). The quality of the clad layer is of importance as it forms the building block of the prototype. The next issue of concern is the bonding (fusion) of layers to build the prototype. Combining these issues together with the460ability of the system to orient the part in the required direction during deposition makes it possible to build parts with complex contours. Experiments were also carried out successfully to develop prototypes based on these principles (Kreutz et al., 1995).

Experiments were also carried out successfully to develop prototypes based on these principles (Kreutz et al., 1995). Laser Engineering Net Shaping (LENS) has demonstrated the feasibility to fabricate geometrically complex shapes with functional materials directly from a CAD model (Keicher et al., 1998). Integration of layered manufacturing and material removal processes is the latest trend in rapid prototyping. This enables the manufacture of complex geometries with increased accuracy and surface finish. The deposition of each layer is followed by machining of the excess material, which is often referred to as “exterior sculpting” (Kulkarni et al., 2000).This paper reviews the successful application of Laser Cladding for RP and the control issues for this application.

Laser cladding process dynamics and process parameters are studied to determine their importance in real-time control of the laser deposition process.Previous Work in Laser Cladding for Rapid PrototypingLaser Cladding is a material processing technique in which a laser is used as a heating source to melt the alloy powder to be cladded onto a substrate. The application of this technique is being extended to obtain layers of deposition of desired height and width with superior properties in terms of pureness, homogeneity and surface finish. Thus, a metal prototype can be generated by selective cladding point–by–point and layer–by–layer. Some efforts have been made to produce metal prototypes by a layer additive approach (Kosh et al., 1993). They have examined building parts in one and two dimensions, taking into consideration the time and cost involved in the process as compared with traditional methods.

They have examined building parts in one and two dimensions, taking into consideration the time and cost involved in the process as compared with traditional methods. Direct Light Fabrication (DLF), which is being developed at Los Almos National laboratory, has proven capable of producing metal parts with reasonably good accuracy and improved metallurgical properties (Lewis et al., 1994; Milewski et al., 1998). It is based on the same principle of adding layers until a near net- shape part is obtained. One such application was carried out using a CO2 laser with inert gas propulsion of material powder into the molten pool generated by laser radiation on the substrate. A specially designed nozzle was used for this specific application. The movement of the substrate in X and Y directions in combination with the movement of the optics in Z direction and a simultaneous change of nozzle angle allowed the generation of arbitrary three–dimensional structures (Kreute et al., 1995).

The geometry of the clad was also investigated by metallographic techniques in combination with optical microscopy. Laser Direct Cladding (LDC) and Selective Laser Cladding (SLC) are based on the same principle as laser cladding, where a gas jet containing fine metal powder is directed via a coaxial nozzle through the path of the laser beam, which is focused slightly above the workpiece. The powder heats up to form molten particles and this laser/melt stream is traversed across the workpiece. A small melt–pool is formed on the surface of the workpiece where the molten particles land to form a layer upon cooling, about 0.5 mm thick, after the laser beam has moved on. It is also possible to lay down very narrow tracks in this way (0.8 mm) and multiple layers can be deposited on top of each other, allowing the formation of complex parts in a relatively short time (Morgan et al., 1997; Jeng et al., 2000).

Laser Engineering Net Shaping (LENS) is one developing rapid metal forming process that has demonstrated the feasibility of laser metal forming to produce near–net shape metal parts. It utilizes STL file rendition of a CAD solid model to build an object one layer at a time. The file is sliced electronically into a series of layers that are subsequently used to generate461the motion to deposit each layer of the material. These layers are then deposited in a subsequent fashion to build the entire part (Keicher et al., 1998). Direct Metal Deposition (DMD) is being used to fabricate molds and dies and for part repair. Its wide application in aerospace and medical fields is due to its large savings in cost. Mazumder et al.

Direct Metal Deposition (DMD) is being used to fabricate molds and dies and for part repair. Its wide application in aerospace and medical fields is due to its large savings in cost. Mazumder et al. (1999) demonstrated the application of laser aided DMD to generate components with dimensional accuracy of 0.01 inch with the required control of process parameters to obtain the desired microstructure.Relevant Parameters and Factors Effecting Laser CladdingThe characteristics of a built up part using the above–mentioned techniques mainly depends on the clad properties. Therefore, the parameters governing the cladding process have to be studied carefully. These parameters play an important role in determining the clad profile, dilution of the cladding metal, fusion between layers, homogeneity of the layers, surface finish, defects such as porosity, cracking due to thermal stresses, plasma formation, etc.

Dilution is an important factor and a desired range should be set to determine various other factors and parameters governing the laser cladding process. It determines the thickness of the liquid layer on the substrate to ensure the bonding of the current layer with the previous layer. It is not possible to predict the influence of an individual parameter on the cladding process. In general, several parameters have to be varied simultaneously to obtain the desired characteristics of the deposited layer. There are also several limitations that restrict the variations of process parameters. Vetter et al. (1994) performed experiments to determine the state of powder when it arrives on the substrate surface. This forms an important limitation as the particles of the powder should be hot enough to adhere to the surface and form a layer, but not too hot such that the particles vaporize, followed by ionization and plasma formation.

Thus, the energy available per unit length of clad pass per unit mass of powder and interaction time form the key factors in controlling the state of the powder, provided the other parameters such as powder mass flow rate, CNC table traverse speed, etc. are kept constant. The Heat Affected Zone (HAZ) induces surface distortion and residual stress and may be critical when producing small parts. In overlapping of tracks, the HAZ structure of prior tracks may be tempered by subsequent passes. The width and depth of the HAZ have to be used as indices for determining process performance in some cases. Once the conditions are set to control the above factors, the process parameters can be varied in a defined range to obtain the required clad height and thickness, which in turn are determined by bead geometry and overlap factor. These parameters depend on the powder feed rate, power density and CNC table feed rate.

These parameters depend on the powder feed rate, power density and CNC table feed rate. Some common surface defects such as the staircase effect (occurs due to part slicing of part), chordal effect (induced when a STL file is generated from the CAD model), support structure burrs and errors due to starting and ending of deposition are to be minimized to eliminate dimensional inaccuracies and improve surface finish. Different control techniques have to be applied to optimize these system parameters to accomplish the required quality and precision in fabricating a part. By exercising on–line control of these parameters, complex metal parts may be built by adding layers. Care should be taken to avoid porosity, which occurs due to cavities between tracks that form from overlapped tracks or the evolution of entrapped gasses in the clad tracks. This can be avoided by proper choice of overlap factor, which also determines the surface finish of the part.

The starred (*) parameters may be varied on–line.State–of–the–Art in Laser Cladding ControlRapid Prototyping via lasers requires synchronization of three basic components of the laser deposition system, Powder Feeder System, Energy Delivery System, and CNC workstation. To enhance the part quality, close monitoring and control of the variables of these systems are required. Feedback controllers have to be designed to regulate these variables mainly to control melt pool size, temperature distribution in the melt pool, cooling rate (for microstructure manipulation) and clad height and width. These variables may vary during the operation due to463fluctuations in system parameters such as powder flow rate, beam position and diameter, output power, and CNC feedrate, and pre–setting these operating parameters is not appropriate. They have to be monitored and optimized continuously to obtain the desired conditions.

They have to be monitored and optimized continuously to obtain the desired conditions. This forms the initial step for real-time process control.The dimensional accuracy of the part depends on the uniformity and repeatability of the clad height and width being deposited. Mazumder et al. (1999) described the application of multiple sensors for closed–loop feedback control of the bead height. The height controller shuts off the laser until it passes the excess built up region, thus preventing the powder from melting. An alternative way to control the bead dimensions is by regulating the powder flowrate, provided the traverse speed of the CNC table is kept constant. Regulation of powder flow rate controls the dilution for a given powder density. Also the carrier and shielding gas flowrate can influence the amount of powder being deposited. By increasing the carrier and shielding gas flowrate, the excess powder can be blown out of the way of the laser beam (Mazumder et al., 1999).

By increasing the carrier and shielding gas flowrate, the excess powder can be blown out of the way of the laser beam (Mazumder et al., 1999). Most of the powder feeder systems used for laser cladding were open loop without flow rate sensing. The basic disadvantage of these were their inability to control the flowrate which continuously changes due to variations in powder volume density, with time and level of powder in the hopper although the control settings are kept constant. A small deviation in mass flowrate results in large variations in the geometry and microstructure of the produced tracks. To account for these problems, Li and Steen (1993) designed a closed–loop control system employing Proportional plus Integral plus Derivative (PID) controller with a feed–forward strategy for on-line feedback control of powder flowrate. Carvalho et al.

Carvalho et al. (1995) designed a closed loop control system using a (PID) controller for independent delivery of two different powders for variable composition laser cladding. A specially designed coaxial nozzle was used which increased the powder utilization efficiency from approximately 30–50% to values higher than 80% (Hu et al., 1997; Lin and Steen, 1998).Temperature is another critical factor that requires continuous monitoring and control. It determines the melt pool dimensions and, hence, the dilution. If the temperature is too low, the resulting melt pool catches little powder and if the temperature is too high, it may melt back the workpiece. Morgan et al. (1997) described an effective means of controlling the temperature by controlling the laser power via positioning the laser focus relative to the workpiece. It also required a constant adjustment as the height of the structure steadily increases.

It also required a constant adjustment as the height of the structure steadily increases. Experiments were performed to demonstrate the effectiveness of closed–loop over open–loop control of these parameters. Li et al. (1987) developed an in–process laser control loop, which is based on an algorithm involving tune currents. The system used a microprocessor based in–process beam control unit using beam sensing via a Laser Beam Analyzer (LBA). Derouet et al. (1997) estimated the melt pool depth from the surface width and maximum temperature of the melt pool. This melt pool depth was controlled by a feedback loop using a PID controller. Laser power and scanning velocity were used as the parameters to control the depth of melt pool. Li et al. (1987) developed a real-time expert system and a laser cladding control system to determine the optimal operating conditions for a given requirement and for online fault diagnosis and correction. Koomsap et al.

Koomsap et al. (2001) presented a simulation–based design of a laser based, free–forming process controller. A simplified model called metamodel was introduced to express the relationship between process characteristics and three process parameters: laser power, traverse speed, and powder feedrate. A dynamic metamodel was obtained and a temperature feedback controller was used to regulate the process. A Proportional plus Integral (PI) controller464was used to regulate the system. Bouhal et al. (1999) and Han and Jafari (1999) proposed a tracking controller for positioning and deposition accuracies in part fabrication for fused deposition processes. Fang and Jafari (1999) designed a statistical feedback control architecture integrating Statistical Process Control (SPC) and Automated Process Control (APC) to adjust parameters such as powder flowrate to minimize the possible defects in the next layer.

They focused on on–line process parameter adjustment using a layer–to–layer controller. Doumanidis and Skoredli (2000) established a dynamic distributed parameter model with in-process parameter identification to generate a 3D surface geometry. Geometric predictions were made by a real-time model. A controller was designed to regulate the part geometry taking advantage of these predictions. Table 1 shows some of the process control techniques discussed in various papers.ReferencesLaser Power (kW)Traverse Speed (mm/sec)Powder Flow Rate (g/min)Process CharacteristicsControlling ParametersControl SystemPrincipleMorgan et al. (1997)Hu et al. (1997)Meriaudeau et al. (1997)Fang et al.

Control techniques applied to processes such as dilution, laser power465distribution, powder flowrate, melt pool depth, etc., and on–line control of the parameters governing these processes were discussed.The study of various works indicates that systematic implementation of process control requires a complete understanding of relation between various parameters and its effect on individual processes and the system as a whole. Most authors have put in efforts in designing closed loop control systems for real-time application of laser cladding for RP. In doing so, many assumptions such as the CNC feed rate and powder flowrate were assumed constant in determining the effect of laser power on processes such as dilution. In real-time application these may not be constant, as the powder flowrate keeps changing as the volume density of powder keeps varying depending on the level of powder in the hopper.

In real-time application these may not be constant, as the powder flowrate keeps changing as the volume density of powder keeps varying depending on the level of powder in the hopper. Also there might be slight variations in the CNC feedrate, due to accelerations and decelerations while depositing at the edges of the prototype being built. These may lead to overfills and underfills. Hence these deviations need to be taken into consideration and feedback control systems have to be designed to account for these deviations and for process automation and control for optimal operating conditions.

Commercial application of laser cladding for RP requires cost oriented design and operating conditions of the laser deposition system, mainly the powder feeder systems, which have to be analyzed to increase powder utilization efficiency which is as low as 30–50% in many of the experimental works.ReferencesBouhal, A., Jafari, M.A., Han, W., and Fang, T., 1999, “Tracking Control and Trajectory Planning in Layered Manufacturing Applications,” IEEE Transaction on Industrial Electronics, Vol. 46, No. 2, pp. 445–451.Carlvalho, P.A., Braz, N., Pontinha, M.M., Ferreira, M.G.S., Steen, W.M., Vilar, R., and Watkin, K.G., 1995, “Automated Workstation for Variable Composition Laser Cladding – Its Use for Rapid Alloy Scanning,” Surface and Coatings Technology, Vol. 72, pp. 62–70.Derouet, H., Sabatier, L., Coste, F., and Fabbro, R., 1997, “Process Control Applied toLaser Surface Remelting,” Proceedings of ICALEO, Sec. C, pp. 85–92.Doumanidis, C.

ntrol of Residual Thermal Stresses in Shape Deposition ManufacturingR.K. Chin, J.L. Beuth, C.H. Amon Department of Mechanical Engineering and Engineering Design Research Center Carnegie Mellon University Pittsburgh, PA 15213-3890AbstractLayer-level thermal and mechanical modeling of the microcasting stage of ShaPe Deposition Manufacturing is presented. Thermo-mechanical models of carbon steel deposited onto a carbon steel substrate are described. Mechanics modeling addresses the issue of residual stress build-up. The effects of substrate heating and bending constraint on the build-up of residual stresses are shown. Results show that thermal cycling from newly applied droplets drastically changes the stress state in the top of the substrate. Originally unstressed regions go through an inelastic compression-tension stress cycle. Residual stresses reach values that may cause yielding in carbon steel.

Originally unstressed regions go through an inelastic compression-tension stress cycle. Residual stresses reach values that may cause yielding in carbon steel. Moderate heating of the substrate above room temperature prior to droplet deposition reduces stresses significantly. Bending constraint during part manufacture allows partial relaxation of stresses as the constraint is removed.IntroductionShape Deposition Manufacturing (SDM) is a manufacturing process which allows the creation of complex three-dimensional parts by the successive deposition of layers of material. After each layer is deposited, it is machined to the required dimensions before depositing the next layer (Merz et al., 1994). SDM combines solid freeform fabrication with other processing operations, such as multi-axis CNC machining, material deposition, and shot peening. The benefits of SDM include the ability to build parts with complicated geometries, multiple materials, and embedded components.

The benefits of SDM include the ability to build parts with complicated geometries, multiple materials, and embedded components. Unlike conventional solid freeform fabrication processes, SDM can also directly build fully dense, functional metal parts to high dimensional accuracy.Within SDM, a process for depositing layers is required. The principal deposition process currently in use is termed microcasting, in which droplets of molten material are deposited onto a much cooler substrate. Residual stresses arise when the hot droplet cools off while it heats up the material underneath, causing differential thermal strains. Undesirable effects of residual stress include warping, loss of edge tolerance, and delamination. Residual stress can also reduce the apparent strength and service life of a part. Residual stress build-up is inherent in any manufacturing process based on successive molten material deposition.

Residual stress can also reduce the apparent strength and service life of a part. Residual stress build-up is inherent in any manufacturing process based on successive molten material deposition. Ultimately, to control the undesirable effects of residual stresses through process changes and part design changes, it is necessary to understand how such stresses build up during manufacture.The goal of this work is to model microcasting on a layer level. The temperature distribution in the molten droplet and the substrate, and the stresses that build up in the droplet and substrate due to differential thermal strains, are investigated after the droplet spreads out until the temperature distribution becomes uniform. A one-dimensional thermal model is first presented and verified. The corresponding mechanical model undergoing thermal loads is verified by using an assumption of pure elastic behavior.

A one-dimensional thermal model is first presented and verified. The corresponding mechanical model undergoing thermal loads is verified by using an assumption of pure elastic behavior. Then a model incorporating creep and temperature-dependent properties is used to predict thermal stresses in microcasted configurations. The solution strategy is similar to that presented by Thomas et al. (1987) and by Zabaras et al. (1991). Finally, the stress reductions achievable from uniform heating of the substrate and constraint of the substrate are shown.221Thermal ModelResidual stress modeling has both a thermal aspect and a mechanical aspect. The approach used here is to solve the thermal problem first and to use then the temperature solution as an input to the mechanical model. In the mechanical model, loading comes from differential thermal expansions and contractions. Mechanically-induced temperature changes are assumed to be negligible.

In the mechanical model, loading comes from differential thermal expansions and contractions. Mechanically-induced temperature changes are assumed to be negligible. The ABAQUS finite element package is used in this study because of its capability to model both the thermal and mechanical problems.The thermal model is shown in Figure 1a. The geometric parameters are the thicknesses of the droplet and substrate. The three initial temperatures of the model are the initial temperatures of the substrate, droplet, and interface. The bottom of the substrate is fixed at the initial substrate temperature. At the top of the substrate, boundary conditions corresponding to heat convection to the ambient and radiation from the plasma arc are imposed.

At the top of the substrate, boundary conditions corresponding to heat convection to the ambient and radiation from the plasma arc are imposed. The temperature-dependent conductivity, specific heat, and diffusivity used in the model (from Allard, 1969, and Touloukian, 1967) are those of a low-carbon steel.To verify the accuracy of the thermal modeling, a typical finite element result is compared to an independent finite difference result from Amon et al. (1995). The specific parameters used by Amon et al. (1995) are shown directly in Figure la, except for the geometry and freezing parameters. The geometry is of a 1 mm layer deposited onto a 7 mm substrate, and the latent heat of fusion is 270,000 Jlkg. In the finite difference model, the material solidifies at exactly 1773 K, while in the finite element model, the material has liquidus and solidus temperatures of 1775 K and 1771 K, chosen to correspond to a mean solidification temperature of 1773 K.

Distinct liquidus and solidus temperatures are used to incorporate latent heat effects into an effective specific heat, and the 4 K difference between the liquidus and solidus temperatures is negligible. The comparison of the finite difference and finite element solutions for the temperature history at a point located 0.5 mm into the droplet (halfway between the top of the substrate and the top of the droplet) yields a maximum discrepancy of 30 K, which corresponds to 1.32% of the working temperature range of 303 K to 2573 K.convection h =5.4 W/m ambientT =2_K 323K,radiation 120 W/m2droplet initial T = 2573 K~r ~initial interfa T =1732 Kfix bottom temperatureceat 303 K,substrate initial T = 303 KFigure la. Thermal Model.Figure lb. Mechanical Model.To match the geometry of a layer and substrate used during part manufacture, a thermal model of a 3 mm layer deposited on top of a 19 mm substrate is constructed.

Mechanical Model.To match the geometry of a layer and substrate used during part manufacture, a thermal model of a 3 mm layer deposited on top of a 19 mm substrate is constructed. It consists of a uniform mesh of 220 one-dimensional first-order elements. The liquidus and solidus temperatures in this model are 1785 K and 1748 K, which are the estimated liquidus and solidus for a low-222carbon steel used in a part made by SDM. The other parameters are identical to those used to verify the thermal modeling. A mesh refinement study using three levels of resolution showed negligible changes in the temperature distribution when the number of elements was doubled or halved.Figure 2 shows the temperature distribution through the layer and substrate at discrete . times. The deposited layer corresponds to the positive axial coordinate while the substrate c:orresponds to t.he negative axial coordinate (see Figure la).

Second-order axisymmetric elements are used in the mechanical model with appropriate boundary conditions applied to render the stress distribution a function of the axial coordinate only. As with the thermal model, a uniform mesh of 220 elements is used, and three levels of mesh refinement were tested to confirm sufficient mesh resolution.The finite domain used for all the mechanical results presented in this work is shown in Figure lb. A condition of zero axial (z) displacement is applied at the bottom of the model. This corresponds to a flat substrate and also prevents rigid-body motion in the axial direction. The body remains free to expand or contract in the axial direction. The centerline undergoes zero radial displacement while the outer wall is constrained to expand (or contract) uniformly in the radial direction.

The centerline undergoes zero radial displacement while the outer wall is constrained to expand (or contract) uniformly in the radial direction. Because this model maintains straight vertical edges and a flat substrate bottom, it allows no bending deformation.To model a deposited layer of indefinite extent, periodic boundary conditions are needed. Periodic boundary conditions require straight edges (but not necessarily vertical edges) on the finite domain to allow any number of the finite domains to fit together as they deform.

Periodic boundary conditions require straight edges (but not necessarily vertical edges) on the finite domain to allow any number of the finite domains to fit together as they deform. While axisymmetric elements are used for the mechanical model, the periodic boundary conditions and223the absence of radial temperature variation (aTfar = 0) lead to a biaxial stress state that can be described with only one stress component (either O"rr or O"ee) that is a function of the z coordinate only.A mechanical solution which assumes linear elastic behavior with temperature-independent mechanical properties (Young's modulus, Poisson's ratio, and thermal expansion coefficient) in the droplet and substrate is used to verify the correct use of ABAQUS to solve the mechanical portion of the problem. The purely elastic material model has a steady-state stress solution that can be expressed in closed form.

The purely elastic material model has a steady-state stress solution that can be expressed in closed form. Also, results from an elastic solution are easy to interpret, so this solution aids the interpretation of subsequent results including nonlinear deformation.Figure 3 shows the stress (O"rr or O"ee normalized) as a function of depth in the deposited layer and substrate at discrete times. There is no stress initially, but as the hot droplet cools, tensile stresses build up in the droplet. Meanwhile, compressive stresses build up in the substrate. There are two causes for compression in the substrate. First, the top of the substrate is restrained from expanding freely as it heats up. Second, the mechanical boundary conditions demand zero net radial force; this condition requires compression in the substrate to balance the tension in the newly deposited layer. A compressive stress remains even after the entire domain cools to room temperature.

A compressive stress remains even after the entire domain cools to room temperature. Because the deposited layer and the substrate undergo equal uniform strains and experience no external forces, the final stresses must be(1v)--'---- =EailT 1+1sand(1 - v)EailT= - - s 1 + d(1)whereO"d =final radial stress in the droplet O"s =final radial stress in the substrate E Young's modulus v Poisson's ratio a =linear thermal expansion coefficientilT = initial temperature difference betweendroplet and substrated = thickness of the droplet = 0.003 m s = thickness of the substrate = 0.019 mThe final normalized residual stress values are predicted to be 0.86 in the droplet and -0.14 in the substrate, which agrees with the finite element result.predict stresses in microcasted parts, a model is formulated that incorporates a creep law, temperature-dependent Young's modulus, and temperature-dependent linear thermal expansion coefficient.

To account for the effect of creep, a secondary (steady-state) creep law for a medium carbon steel in the austenitic phase is chosen (from Thomas et ai., 1987).wheree= A (sinh(Bcr))n exp(-i) E=creep strain rate in reciprocal seconds cr =Mises equivalent stress in MPa T temperature in K A 907 x 1010B == 0.0356 C = 41938 n =6.9(2)This creep law is a commonly used form based on activation energy arguments and is used by Thomas et ai. (1987) to study casting problems. In the creep law (equation [2]), the interaction between the exponential factor and the hyperbolic sine effectively builds in a temperature dependent yield stress with perfect plasticity (no hardening). At a temperature of 303 K, a stress224larger than 460 MPa will cause a high creep strain rate; the effective yield stress at 303 K is 460 MPa.The piecewise-linear temperature-dependent Young's modulus used in the model is from data of Thomas et al.

(1987) and approaches a value of about 25% of the room temperature value at the solidus temperature. At 303 K, the Young's modulus is 207.75 GPa. The temperature- dependent linear thermal expansion coefficient, ranging from 11.9 x lO-6/K to 14.9 x lQ-6/K, is taken from the ASMMetals Reference Book (1981), where it is defined as the mean expansion between 293 K and T, where T ranges from 373 K to 973 K. Beyond that range, a constant value is used. The Poisson's ratio used in the model is a constant 0.27.Figure 4 shows the radial stress distribution obtained at discrete times with temperature dependent creep, Young's modulus, and thermal expansion. The tension in the deposited layer builds up· slower than in the pure elastic case because creep at high temperatures does not permit high stresses. Very early, there is compression in the top of the substrate due to heating just as for the elastic case, although the magnitude is smaller than for the elastic case.

Very early, there is compression in the top of the substrate due to heating just as for the elastic case, although the magnitude is smaller than for the elastic case. As seen in Figure 2, heat conducted from the molten droplet into this region raises the temperature hundreds of degrees Inelastic straining occurs while the material is under compression at these elevated K. temperatures. After the material reaches its maximum temperature (which is different for each point), it begins to cool but is restrained from freely contracting. As this happens, the stress state turns to tension. The final stress distribution (see the solid line in Figure 4) consists of nearly uniform tension in the newly deposited layer, tension at the top of the substrate, and nearly uniform compression at the bottom half of the substrate. The compressive stress at the bottom of the substrate is a reaction to the tensile stresses in other portions of the droplet and substrate.

The cases considered thus far use an initial substrate temperature of 303 K. If a higher substrate temperature is used, then a reduction of residual stress can be expected. Preheating is common in welding applications. A quantitative answer is sought to the question of whether a moderate increase in substrate temperature can significantly reduce residual stresses in microcasting. Even225though substrate heating makes the manufacturing process more complex, .a large benefit. rr:'-ay justify the added complexity. Moderate uniform preheating may be practIcal, and exarmmng uniform preheating also represents the first step toward examining local preheating in the future.Two higher substrate initial temperatures of 473 K (200 "C) and 673 K (400 "C) are used to examine the potential benefits of heating the substrate. 473 K represents a moderate preheating of the substrate while 673 K represents a practical limit.

473 K represents a moderate preheating of the substrate while 673 K represents a practical limit. At high preheating temperatures, problems may be encountered involving oxidation of the part exterior or coarsening of the highly desirable fine-scale microstructure achieved by rapid cooling of deposited layers.The stress distributions for initial substrate temperatures of 473 K and 673 K are similar to the stress distribution for an initial 303 K substrate, except that the stress magnitudes are smaller. The steady-state final stress distributions corresponding to the three substrate temperatures investigated are shown in Figure 5. The 303 K stress distribution in Figure 5 is the same as the steady-state stress distribution from Figure 4. Heating the substrate to 473 K and 673 K reduces the magnitudes of the final stresses by about 150 MPa and 250 MPa, respectively, in the deposited layer and by about 200 MPa and 300 MPa, respectively, at the bottom of the substrate.

- • • ~ -. . . . -• ~ -• • -" , - -I: I: I': : .~ • • - ; ! : : .• • • • • J • • •• • • • » •4002000 Stress (MPa)200400Figure 5. Final stress distribution for 3 different initial substrate temperatures.Substrate ConstraintAlthough displacement constraints are usually associated with higher stresses, it is possible to use constraint in microcasting to actually lower the magnitude of the residual thermal stress. To determine the role of constraint, two loading configurations are examined. The mechanical boundary conditions sketched in the basic model of Figure Ib do not allow the substrate to bend, simulating bolts holding the substrate down against a pallet during part manufacture. Two new loading configurations are investigated and shown in Figures 6a and 6b. In Figure 6a, the model is constrained from bending during manufacturing (as in Figure Ib) but is then released.

Two new loading configurations are investigated and shown in Figures 6a and 6b. In Figure 6a, the model is constrained from bending during manufacturing (as in Figure Ib) but is then released. To simulate releasing (unbolting) the substrate after part manufacture, a moment of the same magnitude as the resultant moment that constrains the part from bending, but of opposite sign, is superimposed. . It is added after the stresses during. manufacture reach steady state. The superimposed moment is made up of a linear stress distribution that adds no net force and elastically deforms each point in the layer and substrate. In Figure 6b, the mechanical model is of a substrate that is permitted to bend during manufacture. The outer wall remains straight while it226displaces but is also allowed to rotate, resulting in no net applied moment. These boundary conditions represent microcasting without constraining the substrate during part manufacture.O'rrdropletdroplet+Figure 6a.

These boundary conditions represent microcasting without constraining the substrate during part manufacture.O'rrdropletdroplet+Figure 6a. Constrain from bending during manufacture,then release.Figure 6b. Allow bending during part manufacture.Figure 7a shows the stress distribution before unbolting, the superimposed stresses, and the stress distribution after unbolting for the problem of Figure 6a. The stress distribution before unbolting is the final stress distribution from Figure 4. The largest superimposed stresses are quite large (exceeding 500 MPa). In some locations, the sign of the stress changes after unbolting. For instance, unbolting changes the tension in the deposited layer to compression and the compression at the bottom 3 mm of the substrate to tension. Also, the peak stress magnitude decreases 200 MPa after unbolting.3oE g fd c: 'E <3 ro ~ -15510' : , ' /:: ··c········· .; ." " .. ~.•......... ; : " . : • .:' ..

At the top of the deposited layer, bolting reduces the stress magnitude by about 120 MPa. Bolting also reduces the stress magnitude at the bottom of the substrate. Furthermore, the difference between the largest tensile stress and the smallest compressive stress is 380 MPa when bending is not allowed versus a larger 505 MPa when227bending is allowed during manufacture. An additional benefit of bolting is the control of bending deformation. If the first layer is deposited without bending constraint, then the substrate and layer warp. When a second layer is deposited, curvature change is added to the warped configuration. Each additional layer causes the incomplete part to warp further. This problem will not appear if the part is restricted from bending during manufacture.

Each additional layer causes the incomplete part to warp further. This problem will not appear if the part is restricted from bending during manufacture. Thus bending constraint from bolting to a stiff pallet is important for control of both residual stress and part warping.SummaryThis study has presented verified thermo-mechanical models for a single layer of carbon steel deposited onto an initially stress-free thick carbon steel substrate, simulating the effects of droplets of a microcasted part being deposited onto a thick substrate. The models show that the deposited layer causes significant substrate heating to a depth of 2 to 3 layer thicknesses. For a substrate that is initially at room temperature, thermal cycling and inelastic flow cause the deposited layer and the portion of the substrate that is significantly heated to reach very high stresses in tension. Preheating the substrate moderately (to 473 K) reduces the stresses significantly (by 200 MPa).

Preheating the substrate moderately (to 473 K) reduces the stresses significantly (by 200 MPa). Bending constraint during manufacturing is important for control of warping and because after manufacturing, the residual stresses are partially relaxed after the bending constraint is removed.AcknowledgementsThe authors gratefully acknowledge financial support by the Office of Naval Research under Grant No. NOOOI4-94-1-0183. This work is also supported by the Engineering Design Research Center, an Engineering Research Center of the National Science Foundation, under Grant No.

NOOOI4-94-1-0183. This work is also supported by the Engineering Design Research Center, an Engineering Research Center of the National Science Foundation, under Grant No. EEC-8943164.ReferencesAllard, S., (ed.), 1969, "Metals Thermal and Mechanical Data," "Metaux donnees thermiques et mecaniques," Tables internationales de constantes selectionnees, 16, Oxford, New York, Pergamon Press.Amon, C.H., Schmaltz, K.S., Merz, R., and Prinz, F.B., 1995, "Numerical and Experimental Investigation of Interface Bonding Via Substrate Remelting of an Impinging Molten Metal Droplet," accepted: Journal ofHeat Transfer.ASM Metals Reference Book: A handbook of data about metals and metalworking, 1981, compiled by The Editorial Staff, Reference Publications, American Society for Metals, pp. 167.Merz, R., Prinz, F.B., Ramaswami, K., Terk, M., and Weiss, L.E., 1994, "Shape Deposition Manufacturing," Proceedings, Solid Freeform Fabrication Symposium, The University of Texas at Austin, pp.

1-8.Thomas, B.G., Samarasekera, I.V., and Brimacombe, J.K., 1987, "Mathematical Modeling of the Thermal Processing of Steel Ingots: Part II. Stress Model," Metallurgical Transactions B, Vol. 18B, pp. 131-147.Touloukian, Y.S. (ed.), 1967, Thermophysical Properties of High Temperature Solid Materials, Macmillan, New York.Zabaras, N., Ruan, Y., and Richmond, 0., 1991, "On the Calculation of Deformations and Stresses During Axially Symmetric Solidification," Journal of Applied Mechanics, Vol. 58, pp. 865-871.228.

However, using reclaimed polyamide 12 powder in SLS results in undesirable part surface finish. Very limited research has been done on the improvement of part surface quality and results barely exist on improving or modifying the surface quality of parts using extremely aged powders (powders held close to the heat-affected zones). Aiming to improve the surface quality, we propose a novel approach for SLS with (extremely) aged polyamide 12 powders. By combining material preparation, powder and part characterizations, and SLS with a customized method of post-heating, we obtain parts with improved surface quality (e.g., reduced roughness and porosities, and eliminated un- sintered particles).

Particularly, parts 3D-printed using the 30%-30%-40% new-aged-extremely aged mixed powders exhibit the smoothest and flattest surface with no unmolten particles and nearly zero porosity.IntroductionSelective laser sintering (SLS) is one widely adopted technique in additive manufacturing (AM) to manufacture high-quality polymeric and metallic components [1-4]. Compared with other SLS materials such as metals and ceramics, polymeric powder materials offer benefits in low processing temperatures, controllable flowability and high corrosion resistance [5-7] in trade of strength. Particularly, as one of the most important semi-crystalline thermoplastic polymer materials, polyamide 12 and its reinforced/filled forms generate SLS parts with superior mechanical properties over general amorphous materials [2].The microstructures of polyamide 12 materials consist of a series of carbon atoms and the amide groups (-NHCO-), and show both amorphous regions and crystalline regions [1, 8].

Due to the open chain ends, the molecular structures of polyamide 12 materials are prone to molecular changes at high temperatures and during laser-material interactions. In particular, post- condensation, chain scission and chain crosslinking reactions form the essential degradation mechanisms of polyamide 12 powders [9, 10], and induce different material properties.The formation of large particles aggregated from small pieces is a predominant property change for reclaimed polyamide 12 powders [1]. These large particles cause deteriorated surface finish with unmolten high-melting-point pieces in the specimens [11]. With multiple times of reuse and repeated expansion/shrinkage in the fabrication cycles, the surface of the reclaimed polyamide 12 powders exhibits increasing cracks and fragments, lowering the part surface quality [12].

J. Guo et al. [16] presented an experimental and analytical study to clarify the interrelations between surface quality and process parameters.Despite the aforementioned literature, it remains not clear how to maximize surface quality when using reclaimed powders. Moreover, there is a lack of understanding in surface characteristics of parts sintered using extremely aged polyamide 12 powders. Held close to or wrapped by the heat-affected zones (HAZs)†, these powders go through severe degradations during the sintering process. We show, however, that such expensive materials can be reused to produce parts with fine surface textures, reduced porosities, and free from unmolten particles. By material preparation, powder and part characterizations and SLS with controlled post-heating, we obtain a series of parts using differently degraded powders and different combinations.

By material preparation, powder and part characterizations and SLS with controlled post-heating, we obtain a series of parts using differently degraded powders and different combinations. Then after surface cleaning, we examine the surface morphologies of these parts and evaluate the characteristics of the surface morphologies. The result is that the undesirable surface finish of parts printed using reclaimed polyamide 12 powders can be drastically improved after using the proposed strategy.Design of the proposed SLS for reclaimed polyamide 12The proposed method involves material preparation (Step 1) and SLS with post-heating (Step 2). More specifically, Step 1 includes powder collection and powder preprocess, and Step 2 covers powder mixing, parameter control, SLS with post-heating and evaluation.

More specifically, Step 1 includes powder collection and powder preprocess, and Step 2 covers powder mixing, parameter control, SLS with post-heating and evaluation. Details of these procedures are explained in the following parts.Material preparation Powder collectionNew polyamide 12 powders were purchased from EOS Corp, and the reclaimed powders were collected from standard SLS processes on an EOS P 390 machine. Polyamide 12 powders with 3 different degradation levels were used in this work: (1) new polyamide 12 powders with no additional heat treatment; (2) aged polyamide 12 powders located far away from the HAZs in the SLS chamber and are reused in industries; (3) extremely aged polyamide 12 powders adjacent to or wrapped by the HAZs that are not reused in industries currently.

Powder preprocessDue to the large particles or part debris caused by the high temperature and laser induced degradation/aging, when recoating with the extremely aged powders, the powder bed is uneven and rough on surface. A sieving process was applied to preprocess the extremely aged powders for an even recoating surface. This process was conducted in a fume hood using a sieve with a mesh size of 200 um. The time for sieving a batch of 500 ml powders is around 2 hours including† Areas close to the laser-material interaction during sintering.646preparation and post cleaning. After sieving, 84%~90% of the reclaimed materials can be collected and well recoated.Proposed SLS and post-heating Powder mixingFour different kinds of powder combinations were used in this work: (i) pure powders, (ii) new and aged powder mixture, (iii) new and extremely aged powder mixture, and (iv) new, aged, and extremely aged powder mixture.

The nominal parameter settings selected in this work were: preheating temperature, 160 °C; scan speed, 3000 mm/s; laser power, 18 W; scan spacing: 0.3 mm; layer thickness, 150 µm. In this work, samples are all printed with 3 layers and using the optimized parameter settings selected to explore the part surface quality improvement. SLS with post-heatingProcess: The SLS machine used in the paper is an in-house built open-configuration research testbed, with features comparable to commercial machines. The proposed SLS and post- heating control apply tailored heating after the core laser-material interaction with optimized processing parameters. The post-heating here keeps parts at the preheating temperature (160 °C) for a controlled time after the sintering process. In this work, we tested 0 second (no post- heating), 60 seconds, 120 seconds, and 300 seconds of post-heating to different specimens to explore the influences of post-heating on part surface morphologies.

Table 1 shows details of the proposed five-stage SLS. In Stage 1, we 3D printed the benchmark part using 100% new powders with no additional post-heating, and the part was used as a reference to evaluate the surface qualities of other parts. In Stage 2, parts were printed using 100% new powders with different post-heating time (60 seconds, 120 seconds, and 300 seconds) to identify the influences of post-heating on part surface qualities when using 100% new powders. In Stage 3, we printed parts with 100% extremely aged powders at different post-heating time. After comparing the surface qualities of parts using reclaimed powders with and without post heating, 300-second post heating appears most effective for reclaimed powders. In Stage 4, parts were 3D printed using different powder mixtures; parts were also 3D printed with and without 300-second post- heating to study the effects of post-heating on part surface quality when using mixed powders.

In Stage 5, we used the mixtures of three differently degraded powders (new, aged, and extremely aged powder mixtures) with and without 300-second post-heating.

The results form the basis to identify the effects of post-heating on part surface morphology.Table 1 Post-heating based SLS process (The mixed powders are in volume percentages)Stages Stage 1Stage 2Objectives To print the benchmark sample To explore the influences of post- heating on part surface morphologiesCombinations 100% new powder 100% new powder 100% new powderPost heating No 60 s 120 s647Stage 3Stage 4Stage 5when using new powdersTo explore the influences of post- heating on part surface morphologies when using reclaimed powdersTo explore the influences of post- heating on part surface morphologies when using mixed powdersTo explore part using new-aged- extremely aged mixed powders100% new powder 100% extremely aged powder 100% extremely aged powder 100% extremely aged powder 100% extremely aged powder 50% new + 50% aged 50% new + 50% aged 70% new + 30% extremely aged 70% new + 30% extremely aged 50% new + 50% extremely aged 50% new + 50% extremely aged 30% new + 70% extremely aged 30% new + 70% extremely aged 10% new + 90% extremely aged 10% new + 90% extremely aged 30% new + 30% aged + 40% extremely 30% new + 30% aged + 40% extremely300 s No 60 s 120 s 300 s No 300 s No 300 s No 300 s No 300 s No 300 s No 300 sEvaluation: After surface cleaning to remove debris on the part surface, we used Scanning Electron Microscope (SEM) to compare surface morphologies between the benchmark part and the evaluated parts in Table 1.

Compared features include particle coalescence performances, part porosity, the number of unmolten particles, surface microstructures and roughness. The details are explained in the next section.Experimental results and discussionsSurface quality improvements of the 3D-printed parts Stage 1: printing the benchmark sampleFigure 1 presents the SEM images of the 3-layer benchmark sample printed using 100% new polyamide 12 powders without additional heat treatment. In Figure 1 (a) - (c), the part exhibits a smooth and flat surface with no unmolten particles. Meanwhile, high porosity is observed from these images, suggesting an insufficient densification. At a high magnification ratio of 10000 (Figure 1d), some fine lamellae or spherulitic regions in an amorphous matrix are observed. The spherulites radiate from the center and grow in a ringed extinction pattern.

The spherulites radiate from the center and grow in a ringed extinction pattern. These surface characteristics of the benchmark part are used as references to evaluate the surface quality of the other parts. Stage 2: the influences of post-heating on part surface morphology when using new powdersCompared with Stage 1 (100% new samples with no post-heating), the parts obtained in Stage 2 (printed using 100% new polyamide 12 powders with 60 seconds, 120 seconds and 300 seconds post-heating) exhibit very similar surface morphologies of smooth and flat surfaces with high porosity and no visible unmolten particles.

From Figure 2a to d, we observe a gradual melting and coalescing process with post-heating time increasing. In Figure 2a (no post-heating), multiple layers of insufficiently melt particles are observed, and every two or more particles form a neck-like bonding due to the reduction of the free surface energies of the particles triggered by high temperature and intense laser-material interaction. With 60 seconds of post- heating (Figure 2b), the particles bonded by the neck-like structures migrate together. The migration becomes stronger with longer post-heating (Figure 2c). Finally, a large unit, a well- consolidated surface forms with little porosity and few unmolten pieces (Figure 2d).Comparing Figure 1 (100% new powders with no post-heating) with Figure 2a (100% extremely aged powders with no post-heating), we observe an obvious difference in surface morphology.

Numerous visible unmolten particles arise and fuse together to form porous structures in the parts 3D printed using extremely aged powders (Figure 2a). The results indicate the existence of high-melting-point pieces in the extremely aged powders.

Given the above, we conclude that the post-heating process helps to improve the surface quality of the printed parts using extremely aged powders by maximizing the coalescence and consolidation, and a well- consolidated surface obtains with 300 seconds post-heating (Figure 2d).649(a) Parts using 100% extremely aged polyamide 12 powders without post-heating(b) Parts using 100% extremely aged polyamide 12 powders with 60 seconds post-heating(c) Parts using 100% extremely aged polyamide 12 powders with 120 seconds post-heating(d) Parts using 100% extremely aged polyamide 12 powders with 300 seconds post-heating Figure 2 Parts printed using 100% extremely aged polyamide 12 powders with and without post-heatingStage 4: the influences of post-heating on part surface morphology when using mixed powders650Figure 3 presents the SEM images of the parts using new-aged and new-extremely-aged polyamide 12 powder mixtures with (i) no post heating and (ii) 300 seconds post-heating.

Figure 3 (a) - (c) shows the SEM images of parts 3D printed using 50%-50% new-aged mixed powders, 70%-30%, and 50%-50% new-extremely-aged mixed powders, respectively. Smooth and flat surfaces with several unmolten particles are observed with no post heating (Figure 3 a (i), b (i), and c (i)), suggesting that good particle coalescence behaviors are achieved when using 50% or more new powders. Meanwhile, the unmolten particles dramatically decrease, and smooth and flat surfaces with high porosity are obtained with 300 seconds post-heating (Figure 3 a (ii), b (ii), and c (ii)). These surfaces are similar to those of parts using 100% new powders, suggesting a significant improvement of surface qualities by our post-heating.Figure 3 (d) show the SEM images of parts 3D printed using 30%-70% new-extremely- aged mixed powders.

With no post-heating, samples exhibit severe deteriorated and distorted surface morphologies with irregular holes or porous and plenty of unmolten particles (Figure 3 d (i)). With 300-second post-heating, the improved surfaces have relatively flat morphologies with several unmolten particles (Figure 3 d (ii)).Figure 3 (e) exhibit the SEM images of parts 3D printed using 10%-90% new-extremely- aged powder mixtures. With no post-heating, the SEM images exhibit little completed surfaces with numerous unmolten particles and multi-layer porous structures (Figure 3 e (i)). These surfaces are similar to the images when using 100% extremely aged powders. With 300-second post-heating, the images exhibit very smooth and flat surfaces with almost no porous and several unmolten particles (Figure 3 e (ii)).

In the combination, 30% new powders and 70% reclaimed powders were used. When using 30% new powders with no post-heating, the parts in Figure 4 (i) have better-coalesced surfaces compared to the part in Figure 3d (i). Because much more extremely aged powders were used in Stage 4 (Figure 3d, 70%) than in Stage 5 (Figure 4, 40%), making it more difficult to fuse the materials. With 300-second post-heating, the part 3D printed using 30%-30%-40% new-aged- extremely-aged mixed powders has the smoothest and flattest surface with no unmolten particles and almost no porosities (Figure 4 (ii)).652Figure 4 Parts using 30%-30%-40% new-aged-extremely-aged mixed polyamide 12 powders with no post-heating and with 300 seconds post-heatingDiscussions Unmolten particles, coalescence, roughness and porosityFigure 5 shows the SEM images of 3D printed parts with no post-heating to compare the number of unmolten particles, coalescence performances and roughness.

In general, parts using 100% new powders have the best coalescence performance, and almost no unmolten particles (Figure 5a). However, parts using more extremely aged powders exhibit worse coalesced surfaces with drastically increased unmolten particles (Figure 5 (b)-(h)). Parts using more new powders have smoother and flatter surfaces, while parts using mixed powders obtain worse surfaces as the percentages of reclaimed powders increase from 30% to 70% (Figure 5 (b)-(e) and Figure 5 (h)). In particular, when using 90%~100% extremely aged powders (Figure 5 (f)- (g)), numerous insufficiently melt particles are observed, and no consolidated surfaces form.Figure 5 Comparisons of unmolten particles, coalescence and roughness of parts using polyamide 12 powders of different combinations without post-heatingFigure 6 presents the SEM images of 3D printed parts with 300 seconds post-heating to compare the number of unmolten particles, coalescence performances and roughness.

Except for the part using 100% new powders, the images in Figure 6 exhibit better surfaces with enhanced coalescence, decreased unmolten particles and improved smoothness with the proposed post- heating. In Figure 6 (a)-(d) and (h), the images of parts printed using 100% new powders, 50%- 50% new-aged mixed powders, 70%-30%, 50%-50% new-extremely-aged mixtures, and 30%-65330%-40% new-aged-extremely-aged mixed powders, respectively, the surfaces are smooth and flat without any unmolten particles. However, in Figure 6 (e)-(g), the images of parts printed using 30%-70%, 10%-90% new-extremely-aged mixtures, and 100% extremely aged powders, respectively, the parts exhibit several unmolten pieces. These results indicate that the parts using a high percentage of new powders obtain improved smooth and flat surfaces with no unmolten particles with post-heating. These surfaces are comparable to those of part using new powders.

These surfaces are comparable to those of part using new powders. The parts using a high percentage of extremely aged powders get improved surfaces with post- heating, but there still exist several unmolten particles.In Figure 6, the parts also display different porosities. In Figure 6 (a) to (c), the images of parts printed using 100% new powders, 50%-50% new-aged mixtures, and 70%-30% new- extremely-aged mixtures, respectively, there are more large size pores than the other parts. In Figure 6 (d), the images of parts printed using 50%-50% new-extremely-aged mixtures, the number and size of pores decrease significantly. In the remaining images of Figure 6 (e)-(h), almost no pores are observed. These results suggest that the parts using more new powders tend to have more pores.

In the remaining images of Figure 6 (e)-(h), almost no pores are observed. These results suggest that the parts using more new powders tend to have more pores. The parts using more extremely aged powders with post-heating display lower porosity due to better consolidation.Figure 6 Comparisons of unmolten particles, coalescence and roughness of parts using polyamide 12 powders of different combinations with 300 seconds post-heatingMicrostructuresMicrostructure is another important variable in the printed parts when using reclaimed powders. Figure 7 a and b present the microstructures of the parts with no post-heating and with 300 seconds post-heating at a magnification ratio of 10000, respectively. Different powders and powder mixtures yield different microstructures.

Parts using new powders show fine lamellae. Extremely aged powders lead to coarse spherulites. And parts using aged powders show intermediate654morphologies. It can be concluded that the microstructures of parts are largely impacted by the aging status of powders rather than post-heating.(a) Polyamide 12 parts without post-heating(b) Polyamide 12 parts with 300 seconds post-heating Figure 7 Microstructure examinations of (a) polyamide 12 parts without post-heating, and (b) polyamide 12 parts with 300 seconds post-heating at a magnification of 10000ConclusionsThis work proposes an SLS with post-heating to improve surface quality of 3D printed parts using reclaimed polyamide 12 powders. The proposed method decreases roughness and porosity of the printed parts, and eliminates unmolten particles. The effects of post-heating on the surface quality using different powder mixtures were studied.

The effects of post-heating on the surface quality using different powder mixtures were studied. In particular, SEM reveals surface features, including the number of unmolten particles, coalescence performances, roughness, porosity and microstructures.The tests suggest that post-heating barely affected the surface quality of parts using 100% new powders. The parts using 100% extremely aged powders with no post-heating exhibit multiple layers of insufficiently melt particles. However, a 300-second post-heating yields a well-consolidated surface with little porosity and a drastically reduced un-molten particles. The unmolten particles disappear on the parts using 50% or more new powders with 300-second post- heating, showing smooth and flat surfaces with high porosity. When using 90% reclaimed powders, numerous visible unmolten particles and multi-layer porous structures occur in the case with no post-heating.

When using 90% reclaimed powders, numerous visible unmolten particles and multi-layer porous structures occur in the case with no post-heating. With the proposed 300-second post-heating, we obtain smooth and flat surfaces with almost zero porosity and only few unmolten particles. The surface morphologies of parts 3D-printed using the 30%-30%-40% new-aged-extremely-aged mixed powders with 300- second post-heating are even better than parts 3D-printed using 100% new powders.AcknowledgmentsThe work was supported in part by a research grant from Unilever and by NSF award1953155.655References1. S. Dadbakhsh, L. Verbelen, O. Verkinderen, D. Strobbe, P. Van Puyvelde, J. Kruth. Effect of PA12 powder reuse on coalescence behaviour and microstructure of SLS parts. European Polymer Journal 2017;92:250-262. 2. G. Wang, P. Wang, Z. Zhen, W. Zhang, J. Ji. Preparation of PA12 microspheres with tunable morphology and size for use in SLS processing. Mater Des 2015;87:656-662. 3. P.

ntrol of the Cross Section Geometry of Extruded Dental Porcelain Slurries for Rapid Prototyping ApplicationsJiwen Wang, Leon L. Shaw, Harris L. Marcus Department of Metallurgy and Materials Engineering Institute of Materials Science University of Connecticut, Storrs, CT 06269T. B. Cameron Dentsply Ceramco Burlington, NJ 08016AbstractThis study investigates the dependence of the cross section geometry of extruded dental porcelain slurries on the rheological property of the slurry and the extrusion conditions. It is found that a pseudoplastic slurry is a basic requirement for obtaining extruded lines with rectangular cross sections. The cross section geometry of the extrudate is also strongly affected by extrusion parameters including the extrusion nozzle height, nozzle moving speed, and extrusion rate. Proper combinations of these extrusion parameters are necessary in order to obtain extrudates with near rectangular cross sections.

Proper combinations of these extrusion parameters are necessary in order to obtain extrudates with near rectangular cross sections. The results obtained have been explained in terms of the interactions among the rheological properties of the slurry, the shear rate imposed on the slurry during extrusion, the wettability of the slurry on the substrate, and the forced flow of the slurry during extrusion.Keywords: Slurry extrusion, dental porcelain, rheology of slurry, solid freeform fabricationI. Introduction3D objects fabricated through line-by-line and layer-by-layer approaches require that each line deposited has a rectangular cross section in ideal cases. Lines with such cross section geometry can prevent or minimize the formation of voids between lines and layers.

Lines with such cross section geometry can prevent or minimize the formation of voids between lines and layers. This study investigates the dependence of the cross section geometry of extruded dental porcelain slurries on the rheological property of the slurry and the extrusion conditions for dental restorations via the multi-material laser densification (MMLD) process [1-3].Currently, there are two general routes for preparing pastes and slurries for SFF. One requires the use of polymers or resins as the binder in order to maintain the shape of the extrudate [4,5]. These polymers and resins are burned out in the subsequent heat treatment while contamination may remain in the extrudate. The second route is to prepare aqueous pastes or3921slurries without addition of binders and to control the cross section geometry of the extrudate by adjusting the pH value, salt concentration, and volume fraction of solid in the paste [6-11].

Using fine Al2O3 powder, Cesarano, et al [7] have shown that when a slurry is pseudoplastic and drying rates are appropriate, the lines extruded yield nearly rectangular cross sections with relatively straight walls and flattened tops. The paste formulation and particle size have also been shown by Du, et al [11] to have strong effects on the extrudability of thin wall thickness tubes and their thickness and uniformity.Although the effects of rheological behavior of pastes and slurries on the cross section geometry of the extrudate have been studied by several researchers [6-11], the dependency of the cross section geometry on extrusion parameters has never been systematically investigated. To address this issue, the effects of extrusion parameters on the cross section geometry of the extrudate in conjunction with the effects of rheological behavior of dental porcelain slurries have been investigated in this study.

The extrusion parameters investigated include the nozzle height of the extruder, nozzle moving speed, and extrusion rate, all of which are experimental variables and thus can be adjusted to achieve the desired cross section geometry. The results indicate that the effects of these extrusion parameters on the cross section geometry of the extrudate can be explained in terms of the interactions among the rheological properties of the slurry, the shear rate imposed on the slurry during extrusion, the wettability of the slurry on the substrate, and the forced flow of the slurry during extrusion.II. ExperimentalThe dental porcelain powder was provided by Degussa-Ney Dental Inc., Bloomfield, CT. The chemical composition of the porcelain is confidential; however, it is within 5% of the nominal composition of the Weinstein patent [12], which has the following composition (wt%): 63.40% SiO2, 16.70% Al2O3, 1.50% CaO, 0.80% MgO, 3.41% Na2O, and 14.19% K2O.

The as- received powder has angular shapes and their equivalent particle sizes range from 1 to 50 micrometers. To reduce the particle size of the porcelain powder to sub-micrometers, ball- milling process was conducted using a single axis mixer machine loaded with Al2O3 balls. After milling, the powder was dispersed in de-ionized water and the submicron-sized particles were separated from coarse particles (> 1 mm) through sedimentation. Only the submicron-sized particles were used to prepare the slurries.The porcelain slurries were prepared using de-ionized water as the solvent with a solid loading of 40 – 50 vol %. Uniform and stable slurries were obtained by gently mixing the slurry for 24 hours using the mixer machine loaded with Al2O3 balls. The porcelain slurry was extruded using a MMLD machine designed and constructed at the University of Connecticut.

The porcelain slurry was extruded using a MMLD machine designed and constructed at the University of Connecticut. The machine consists of four major components: (i) the process chamber that can have a vacuum down to 10-2 torr, (ii) micro-extruders for powder slurry delivery of up to 3 different materials, (iii) a laser heating system for powder densification with temperature sensing and control system, and (iv) a X-Y-Z positioning system. The pressure exerted on the porcelain slurry within the micro- extruder was applied via an electric cylinder (model: NV-BN23-105B). The slurries were extruded onto SiC plates with a 700mm nozzle. The motion and position control was provided by a computer through a Galil DMC-1800 multi-axis motion control card.

The slurries were extruded onto SiC plates with a 700mm nozzle. The motion and position control was provided by a computer through a Galil DMC-1800 multi-axis motion control card. The extrusion variables3932that could be controlled independently and investigated in this study were the nozzle height of the micro-extruder (i.e., the distance between the nozzle tip and the substrate), the nozzle moving speed (with respect to the substrate, mm/s), and the extrusion rate (i.e., the volume of the slurry extruded per unit time, ml/s).The rheology of the porcelain slurries was measured using a Brookfield DV II digital viscometer. The zeta potential of the slurries was measured at the National Science Foundation Engineering Research Center for Particle Science and Technology in the University of Florida using a Zeta-Reader instrument. The pH value of the slurry was measured using a Denver basic pH meter.

The pH value of the slurry was measured using a Denver basic pH meter. The cross section geometry was defined by the line width and height of the extrudate and its contact angle with the substrate. To measure these geometrical parameters, the porcelain lines extruded were encapsulated using epoxy before cutting and polishing after which these parameters were observed and measured using an optical microscope.III. Results and Discussion3.1 Rheological Behavior of Dental Porcelain SlurriesFigure 1 shows the zeta potential of the dental porcelain powder as a function of the pH value. It indicates that the iso-electric point (i.e.p.) of the dental porcelain powder is at about 5.0, which is between the i.e.p. of Al2O3 (8.7) and SiO2 (2.0). Since Al2O3 and SiO2 are the main components of the dental porcelain [12], the zeta-potential of the porcelain powder appears to be controlled by a combined interactive effect of these two compounds.

Since Al2O3 and SiO2 are the main components of the dental porcelain [12], the zeta-potential of the porcelain powder appears to be controlled by a combined interactive effect of these two compounds. It can be inferred from Figure 1 that slurries with pH < 3.0 or pH > 8.0 should be relatively stable. This is the case because the surfaces of powder particles with these pH values are mainly positively charged (for pH < 3.0) or negatively charged (for pH > 8.0) [13,14]; the same charges on their surfaces provide large repulsive forces between particles, thereby preventing coagulation and thus leading to stable slurries. In contrast, slurries with pH values between 3.0 and 6.5 would be unstable because coagulation of particles will happen quickly owing to the zero or near-zero net charge on the surfaces of these particles.8060) v m40(l a i t n e t o p a t e Z200204012345678910 1112Figure 1.

Zeta potential of the dental porcelain powder as a function of the pH value.60pH value3943Based on Figure 1, clearly, a pseudoplastic slurry with moderate viscosity can be prepared with a pH value near 7.0 – 7.5 because at pH = 5.0 the slurry becomes unstable with high viscosity and at pH > 8.0 or < 3.0 the slurry is very stable. Slurries with pseudoplastic (shear thinning) behavior are highly desirable for controlling the cross section geometry of the extruded line because slurries with shear thinning properties can be extruded with a relatively low extrusion force and solidify in place once the slurry leaves the nozzle due to the removal of shear stresses. Thus, most of the extrusion studies presented below are carried out using dental porcelain slurries with pH = 7.0 – 7.5 unless otherwise mentioned.Figure 2 compares the cross section geometry of the extrudates with pH values equal to 9.3, 8.0 and 7.0.

At pH = 9.3, the slurry is very stable with a relative low viscosity. As a result, the slurry flows continuously after it leaves the nozzle and thus takes on an arch shape with a low contact angle of 40 degrees. At pH = 8.0, the viscosity of the slurry increases slightly and thus the slurry spreads out less on the surface of the substrate, resulting in an increased contact angle of 60 degrees. At pH = 7.0 (approaching the i.e.p.), the slurry becomes pseudoplastic because the low electrical charges of the same sign of the particles allows them to approach to each other, but at the same time prevents them to coagulate. As a result of its shear thinning behavior, the contact angle of the extrudate with the substrate increases to 95 degrees, clearly indicating nearly no spreading after the slurry leaves the nozzle of the extruder.

With a contact angle near 90 degrees, the cross section of the extrudate is approaching a rectangular shape which is highly desirable as a “building block” for fabrication of 3-dimensional objects through layer-by-layer approaches.500 mmpH = 9.3 pH = 8.0 pH = 7.0Contact angle (degree)40 60 95Figure 2. Cross section photos of extrudates with different pH values.3.2 Effects of Nozzle Height on the Cross Section Geometry of ExtrudatesThe distance between the nozzle tip of the extruder and the substrate, termed as the nozzle height hereafter, greatly influences the cross section geometry of the extrudate. The distance between the nozzle and the substrate will limit the space within which the slurry can flow.

The distance between the nozzle and the substrate will limit the space within which the slurry can flow. For a certain slurry extrusion rate and nozzle moving speed, there is a critical nozzle height (hc) above which the cross section geometry of an extruded single line is mainly controlled by the rheological behavior of the slurry and the wettability of the slurry on the substrate. When the nozzle height is lower than hc, the volume of the slurry extruded will be too large for the space between the nozzle and the substrate. As a result, the slurry is forced to spread in the directions perpendicular to the deposited line along the surface of the substrate, and the resultant shape of the extrudate is not only determined by the rheological properties and wettability of the slurry,3954but also by the space between the nozzle and the substrate.

In contrast, when the nozzle height is larger than hc, there is enough space for the deposited slurry so that the rheological properties and wettability of the slurry will determine the cross section geometry of the extrudate.Our experiments indicate that the critical nozzle height can be estimated using thefollowing equation:h = cV d Dv nn(1)where Vd is the volume of the slurry extruded per unit time (called the extrusion rate hereafter), vn the nozzle moving speed, and Dn the nozzle diameter. The physical meaning of eq. (1) is that the volume of the slurry extruded per unit time is equal to the volume available per unit time between the nozzle and the substrate.

The physical meaning of eq. (1) is that the volume of the slurry extruded per unit time is equal to the volume available per unit time between the nozzle and the substrate. When the nozzle height is lower than hc, the slurry is forced to take up the space beyond the volume defined by the product of the nozzle height, the nozzle diameter, and the distance traveled by the nozzle per unit time.Wax500 mmExtrudateSubstrateNozzle height (mm): 200330 670 1000Contact angle (degree): 4065 85 95Figure 3. Cross section of the single lines extruded using a dental porcelain slurry of pH = 7.5 with Vd = 2.5 x 10-3 ml/s, Dn = 0.7 mm, vn = 4.25 mm/s, and different nozzle heights as indicated.Figure 3 shows the cross section of the extrudates deposited with different nozzle heights. The contact angles of these extrudates with the SiC substrate are measured and presented in Figure 4.

The contact angles of these extrudates with the SiC substrate are measured and presented in Figure 4. When the nozzle height equals to 200 and 330 mm, the slurries spread widely along the substrate with low contact angles, which indicates that the nozzle height is lower than hc because the slurry has been forced to flow beyond the space defined by the nozzle height, the nozzle diameter, and the distance traveled by the nozzle per unit time. When the nozzle height equals to 670 mm, the contact angle closes to 90 degrees and no spreading is observed. A further increase in the nozzle height (1000 mm) does not alter the shape of the extrudate much with only a small increase in the contact angle. The critical nozzle height calculated using eq. (1) for the extrusion condition shown in Figure 3 is 840 mm which is between 670 mm and 1000 mm tested in the experiment.

The critical nozzle height calculated using eq. (1) for the extrusion condition shown in Figure 3 is 840 mm which is between 670 mm and 1000 mm tested in the experiment. These results indicate that when the nozzle height is near the critical nozzle height, the contact angle of the slurry on the SiC substrate used in this study is about 90 degrees.3965However, when the nozzle height is substantially smaller than hc (e.g., the cases of 200 and 330 mm in Figure 3), the forced flow is present and the contact angles will be smaller than 90 degrees.100) e e r g e d ( e l g n A90 8070 6050 40t c a t n o C30 2010 0020040060080010001200Nozzle Height(microns)Figure 4.

The contact angle as a function of the nozzle height for slurries with pH = 7.5.3.3 Influence of the Shear Rate on the Cross Section Geometry of ExtrudatesBecause of the pseudoplastic behavior of the slurry, extrusion conducted with high shear rates will have relative low viscosity, whereas extrusion with low shear rates will have high viscosity. In order to investigate the true effects of the shear rate on the cross section geometry of the extrudate, a set of the experiments have been designed in which both the extrusion rate and the nozzle moving speed have been increased proportionally in order to keep the critical nozzle height [see eq. (1)] and the deposition density constant. The deposition density, Dd, is defined as the volume of the slurry extruded per unit time divided by the product of the nozzle diameter and the moving speed of the nozzle.

The deposition density, Dd, is defined as the volume of the slurry extruded per unit time divided by the product of the nozzle diameter and the moving speed of the nozzle. When both the critical height and the deposition density are kept constant, the change in the extrusion rate provides a genuine evaluation of the effect of the shear rate.The results from this special set of the experiments are shown in Figure 5. At high extrusion rates (0.01 ml/s and 0.005 ml/s), the contact angles are smaller than 90 degrees. In contrast, at low extrusion rates (0.0025 ml/s and 0.00125 ml/s), contact angles are greater than 90 degrees. This is so because high extrusion rates result in high shear rates which in turn leads to low viscosities. Thus, when high extrusion rates are used, the slurry flows fast after it leaves the nozzle tip. As a result of the fast flow, the slurry spreads out along the surface of the substrate before it freezes and a low contact angle results.

As a result of the fast flow, the slurry spreads out along the surface of the substrate before it freezes and a low contact angle results. In contrast, when low extrusion rates are used, the flow of the slurry after it leaves the nozzle tip is limited because of its high viscosity. As a consequence of this limited flow, a high contact angle results.It is noted, however, that in spite of the dependency of the cross section geometry on the shear rate, the effect of the shear rate on the contact angle is relatively small in comparison with that of the pH value and the nozzle height.

As seen from Figure 5, the variation of the contact3976angle due to the different shear rates investigated is within 20%, which is much smaller than 100% changes achieved by altering the nozzle height (Figures 4 and 5).500 mmExtrusion rate (ml/s): 0.01 Nozzle moving speed (mm/s): hc (mm): Deposition density (mm3/mm2): 0.84 Contact angle (degree): 85 90 95 1050.005 0.00250.0012517.5 0.848.5 0.844.25 0.842.125 0.840.840.840.84Figure 5: Cross section of the single lines extruded with different extrusion rates and nozzle moving speeds as indicated, while keeping all other processing parameters constant.IV. Concluding RemarksThe present extrusion study clearly shows that in order to get favorable cross section geometry of the porcelain powder line in SFF processes, extrusion parameters including the nozzle height, nozzle moving speed and the extrusion rate should be optimized. However, optimization can be achieved only when the slurry is pseudoplastic.

However, optimization can be achieved only when the slurry is pseudoplastic. The latter can be accomplished by adjusting the pH value of the slurry for the dental porcelain powder to between 7.0 and 7.5. Once the proper slurry is prepared, extrusion conditions can have strong impact on the cross section geometry of the extrudate. There is a critical nozzle height (hc) for a given set of the extrusion rate, nozzle moving speed, and the nozzle diameter. When extrusion is carried out with the nozzle height above the critical nozzle height, the cross section geometry of the extrudate is mainly determined by the rheological behavior of the slurry and the wettability of the slurry on the substrate. When the nozzle height is lower than the critical nozzle height, the resultant shape of the extrudate is not only controlled by the rheological properties and the wettability of the slurry, but also by the forced flow of the slurry along the surface of the substrate.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceControl of Track Morphology in Digital Glass FormingNicholas Cappsa, Jonathan Goldsteinb, Robert Landersa, Edward Kinzela, *aAerospace and Mechanical Engineering, University of Notre Dame, Notre Dame, IN, 46556 bAir Force Research Laboratory, Wright-Patterson AFB, OH 45433 *ekinzel@nd.eduABSTRACT Digital Glass Forming involves locally heating a glass filament using a laser while continuously feeding relative to the workpiece. This enables precise control of the viscosity of the molten region. This viscous melt pool is deformed by forces from the filament and workpiece. This paper investigates the relationships between the input parameters, laser power, feed rate, and scan speed, with the resulting track morphology. Tracks are used to build larger, more complex structures.

Tracks are used to build larger, more complex structures. By underfeeding the filament relative to the able scan speed, the glass is locally drawn to produce tracks with significantly smaller cross-sectional areas than the feedstock material. This technique allows for the production of arbitrary geometries and surface features using a smaller equivalent diameter than the input feedstock. The paper explores the available process zone for this underfed deposition mode and the limitations of this method. Keywords: Additive manufacturing, Glass1.IntroductionSilicate-based glasses are attractive materials for structural, scientific, and optical components due to the material’s unique combination of material properties.

These properties include transparency from the near UV to near IR wavelengths, high thermal stability at elevated temperatures, stiffness, scratch resistance and formability at high temperatures, and the ability to dope the material to change its optical and chemical properties [1-2]. There has been considerable interest in utilizing additive manufacturing (AM) techniques to create optical and structural components from silica-based glasses. However, the material’s high molten viscosity and transparency to common laser wavelengths pose a challenge to most common AM methods. There are significant difficulties with bubble evacuation within powder-based approaches, leading to mechanically weak, low transmission parts [3]. There have been some successful methods for creating transparent glass structures, such as a composite method using silica precursor-loaded polymer slurries deposited using conventional approaches such as direct ink writing or stereolithography.

These parts are then dried, pyrolyzed and densified following ceramic processing techniques to produce dense 3D glass structures. However, there are challenges to utilizing this method for creating parts with high aspect ratios due to limited build volume and part shrinkage, and the method also suffers from low cycle time [4-6]. Other methods use a traditional fused filament fabrication (FFF) approach to deposit glasses but utilize specialized tooling or alternative feedstock chemical compositions (e.g., chalcogenide) to overcome the high temperature requirements for glass deposition [7-8].Another glass AM methodology utilizes a CO2 laser to locally heat solid silica glass feedstocks to deposit the material directly onto a substrate of the same material [9-11]. This method takes inspiration from traditional artisan glass working through the localized application of heating and force to form the feedstock into a bead track, building into the desired shape.

This method takes inspiration from traditional artisan glass working through the localized application of heating and force to form the feedstock into a bead track, building into the desired shape. By articulating the workpiece using CNC-controlled stages, the mechanical freedom of the artisan can be replicated while significantly improving the resolution, leading to the development of the Digital Glass Forming (DGF) process. Using this method, complex 3D artifacts can be created, ranging from fully dense optical forms to sparse free-standing structures, including lattices, using feedstocks ranging from 3 mm to 125µm OD [11 - 13]. However, the piece's form accuracy is limited by this feedstock material's input diameter. While using a 125µm feedstock will allow for high spatial resolution, the surface resolution required for producing optics is significantly smaller than this [14].

While using a 125µm feedstock will allow for high spatial resolution, the surface resolution required for producing optics is significantly smaller than this [14]. This paper aims to explore a technique of underfeeding the filament relative to the table scan rate to locally draw the filament to a smaller equivalent diameter.2. Experimental SetupFigure 1 depicts the DGF setup used in this paper. As shown, 2 mm OD borosilicate feedstock (Duran) is fed into the intersection between a CO2 laser beam (Iradion 1625, λ0=10.6 µm) and a borosilicate substrate. Silicate glasses are highly absorptive at 10.6 µm, so the laser power (P) is approximated as surface a heat flux applied to both the feedstock and the illuminated area on the substrate [15]. For all experiments in this paper, the laser spot size976was selected to be 2 mm FWHM. The substrate is mounted to a four-axis positioning system using computer- controlled stages (Aerotech ANT130XY, ATZ100, and ANT130R).

The substrate is mounted to a four-axis positioning system using computer- controlled stages (Aerotech ANT130XY, ATZ100, and ANT130R). The filament is deposited onto the substrate by feeding in the feedstock at feed rate f while translating the work piece at rate v. The filament is oriented at θF = 40° from the substrate normal, while the laser is incident at θL = 60°. The system is aligned such that the lines that define the filament's front surface, the substrate's top surface, and the center of the laser spot intersect at the same point. This orientation is selected such that the laser energy is directly delivered to the interface between the filament and substrate, increasing the linear deposition rates of glass and allowing for control of the resulting track morphology via loading from the filament feeder, which will be discussed below [16]. Due to the large thermal gradients generated in the process, deposited pieces are subject to residual stress upon cooling.

Due to the large thermal gradients generated in the process, deposited pieces are subject to residual stress upon cooling. To prevent thermal shock and cracking of the substrate, the built plate is maintained at 550°C (above the annealing temperature for borosilicate glass) [17].θFθLω = 2 mm FWHMFigure 1: Photograph of the DGF deposition methodReducing glass feedstock diameters is a common procedure for glass materials, highlighting the unique features of working with glass. While most processes for reducing feedstock diameters to very small equivalent diameters involve pushing molten material through an extrusion nozzle, due to molten glass’ very high molten viscosity and temperature requirements, this method is limited to producing filaments diameters ≥ 2 mm [18].

The most common method for producing glass fibers is drawing the material down from a larger feedstock by continuously pulling on the molten glass, using the viscosity and surface tension to maintain the circular morphology [19]. Taking inspiration from the fiber-pulling techniques, we can underfeed the feedstock material (f) relative to the table scan rate (v), pulling on the filament and locally drawing the equivalent diameter down. By conservation of volume, the equivalent radius change of the track compared to the feedstock is related to the ratio of these two velocities by R/R0 = (f/v)2. For this paper, we arbitrarily chose that the target would be to reduce the 2 mm feedstock diameter down to an equivalent diameter of 0.5 mm, resulting in a target f/v=1/16.

For this paper, we arbitrarily chose that the target would be to reduce the 2 mm feedstock diameter down to an equivalent diameter of 0.5 mm, resulting in a target f/v=1/16. There are diminishing returns on f/v shift to equivalent diameter change beyond this point.Initial experiments with this aggressive underfeeding target showed the most prominent failure mode appeared as the filaments lifting off the substrate, bending in open air. This failure mode can be understood conceptually by examining a simplified force balance model shown in Fig 2 below. In the nominal case at f/v=1 (Left), the primary force delivered to the melt pool is from the filament feeder delivered by the relatively cool back surface of the filament. This is resisted by the normal force of the substrate as well as the viscoelastic deformation of the rest of the filament.

This is resisted by the normal force of the substrate as well as the viscoelastic deformation of the rest of the filament. Additionally, there are more difficult to define forces such as those supplied by the table scanning as well as surface tension between the filament and substrate. As the table scan rate increases and the filament lifts off the substrate, several of these forces disappear (surface tension and substrate normal force). Instead, the force applied by the stage movement is significantly higher along with a significantly higher viscoelastic deformation component as the filament stretches to the desired length to achieve f/v. Developing a mathematical model to describe these forces and their application is a significant challenge, so an effort was made to empirically define the relationships between the input parameters and resulting morphology.

As such an experiment plan was developed to examine the limitations of this approach and identify how close to the f/v target could be achieved.97735 W, f = 0.5 mm/s f/v = 1 Grayscale55 W, f = 0.5 mm/s f/v = 1/8 GrayscaleFTableFFeederFViscousFFeederFTableFSurface TensionFViscousFNormalFigure 2: Comparison of force balance of normal and underfed depositionsA parameter sweep was set up to explore the process zone. For this experiment, a static feed rate f and laser power were set for each experiment. Table scan speed v was then swept to achieve the desired f/v ratio. To increase the data throughput per experiment, the feed rate was varied linearly along the length of the deposition of a straight line, as shown in Fig 3 below. The sample was then sectioned along the dashed red lines such that the cross-section at the end of each section could be viewed.

The sample was then sectioned along the dashed red lines such that the cross-section at the end of each section could be viewed. For this experiment set, the valid track morphologies were divided into two groups depending on the contact angle (θc). To produce fully dense geometries θc ≤90° (termed as slumped) are desired, while tracks that are attached to the substrate but would result in inter-track voids θc ≥ 90° are not. The filled-out process zones for this experiment are shown in Fig 4 below. As shown, in this configuration, the process deposition region is limited to a minimum f/v ≥ ~0.2, due to the presence of this liftoff mode. Notably, the feed rate that maintained the largest viable process zone is at f = 0.5 mm/s, while both higher and lower feed rates had significantly smaller process zones.

Notably, the feed rate that maintained the largest viable process zone is at f = 0.5 mm/s, while both higher and lower feed rates had significantly smaller process zones. The effect of feed rate on the process will be discussed in a later section.f/v=1 f/v= 0.75 f/v= 0.510 mm 10 mm 10 mmf/v= 0.2515 mmTacked (θc>90°) 1 mmθcSlumped (θc<90°)Printing Direction (Reduced f/v)Figure 3: (Left) Diagram of the experimental method, where the track f/v ratio decreases linearly from left to right in defined increments.

(Right) example cross-sections of filaments based on the resulting contact angle.9788.005.334.002.672.001.601.331.141.00253545Power [W]f/v Ratio [1:x]f=0.25 mm/s, ∆z=0 mm8.005.334.002.672.001.601.331.141.0025354555Power [W]f/v Ratio [1:x]f=0.5 mm/s, ∆z=0 mmSlumpedLiftoffTackedBroken8.005.334.002.672.001.601.331.141.0035455565Power [W]f/v Ratio [1:x]f=1 mm/s, ∆z=0 mmFigure 4: Process maps relating the input f/v ratio and input laser power to the resulting track morphology for a few filament feed rates.3.Incorporating shift in ∆zObserving this liftoff failure mode further, it appears that once the filament disconnects from the substrate, the deposition will stabilize printing horizontally at some distance above the substrate zError, as shown in Fig 5 below. This prompted an expansion of the process zone by including a new printing parameter ∆z, representing the upward shift in the substrate to correct for this height error.

This prompted an expansion of the process zone by including a new printing parameter ∆z, representing the upward shift in the substrate to correct for this height error. This was done empirically by repeating the same process zone above (f = 0.5 mm/s), sweeping ∆z and f/v over a few powers. The process zones for these experiments, shown below in Fig 6, show a significant increase in the viable f/v ratios available, with the higher power process zones reaching the desired ratio of 1/16. It is worth noting that at very high laser powers (and sufficiently slow scan speeds), the surface of the deposited glass can overheat, vaporizing important interstitial ions and altering the material properties. At f/v = 1/16, vaporization is minimized due to the increased scan rate. However, this effect places an upper limit on the input power that can be used for depositions.

However, this effect places an upper limit on the input power that can be used for depositions. Additionally, at the highest laser powers at slower feed rates, the filament can get so hot that the molten connection to the substrate becomes unstable, resulting in a disconnection between the materials where the filament balls up (noted in Fig 6).979zErrorFigure 5: Photograph of the ‘liftoff’ failure mode, highlighting the static deposition height off the substrate.16.008.005.334.002.672.001.331.00012345Δz [mm]f/v Ratio [1/x]f=0.5 mm/s, P=55 W16.008.005.334.002.672.001.331.00012345Δz [mm]f/v Ratio [1/x]f=0.5 mm/s, P=65 W16.008.005.334.002.672.001.331.00012345Δz [mm]f/v Ratio [1/x]f=0.5 mm/s, P=75 WBallingFigure 6: Results of the expanded process zone, incorporating ∆z.The significant increase in the viable f/v ratios can be explained by examining how the shift in ∆z affects the temperature and loading profiles of the bead.

Conceptually, the deposition behavior of the glass is determined by the temperature of the interface between the filament and substrate (illuminated by the laser), as well as the temperature on the opposite side of the filament. In short, the temperature at the interface must be sufficiently high to allow the filament and substrate to coalesce into a single piece of material. Additionally, the back surface must be warm enough to allow the filament to bend but still be cool enough to apply force from the feeder to the melt pool. To capture the temperature in these two areas, an infrared camera (Optris PI 640) was used in the directions indicated by Fig 7a below. As shown, when measuring the interface temperature, the maximum temperature of the viewing area is extracted, while on the back surface, an average temperature of a 5 px 5 px area is extracted.

As shown, when measuring the interface temperature, the maximum temperature of the viewing area is extracted, while on the back surface, an average temperature of a 5 px 5 px area is extracted. The smaller defined area on the back surface of the filament is used to distinguish the bending point from the rest of the image, compared to the interface temperature, which is the hottest point of the deposition. Figure 7b shows the effect of980changes in f/v on the temperature profile as solid lines (∆z=0). These data points were repeated for a shift of ∆z=2 mm, shown as dashed lines in the image. As shown in the left figure, there is very little temperature shift on the front surface by changing ∆z (much more sensitive to P and f/v). However, on the back surface, there is a significant decrease in the back surface temperature when increasing ∆z.

However, on the back surface, there is a significant decrease in the back surface temperature when increasing ∆z. This decreased back surface temperature may explain why depositing at a shifted z-plane allows for higher f/v ratios, as the cooler (and therefore stiffer) back surface can re-apply the force from the filament feeder onto the melt pool.TC(A)TBack= 5  5 px averageTCTMax= TFrontΔz=0Δz=2 mm 55 WΔz=0Δz=0Δz=2 mmΔz=2 mm 45 W65 W123456789101112131415161000120014001600180020002200240026002800Vaporization Temp.Front Surface Temperature [°C]f/v Ratio [1/x](B))Δz=2 mm 45 WΔz=0Δz=2 mm 55 W65 WΔz=0Δz=2 mm123456789101112131415166008001000120014001600Back Surface Temperature [°C]f/v Ratio [1/x]Δz=0(C))Figure 7: (A) Diagram of how the front and back surface temperatures were captured with the thermal camera. (B-C) Front and back surface temperatures are plotted as a function of f/v ratio, laser power, and ∆z, respectively.4.

(B-C) Front and back surface temperatures are plotted as a function of f/v ratio, laser power, and ∆z, respectively.4. Effect of Feed RateExpanding the process zone to include other feed rates can significantly alter the viable process zone and deposition mechanics. For instance, the data set for the f = 0.25 mm/s (See Fig. 8 below) has a significantly larger portion of the process zone fail via balling the filament. Heat has more time to diffuse through the filament cross- section at these lower feed rates, resulting in a much more uniform temperature profile. This much hotter back surface temperature reduces the force the feeder can deliver to the melt pool.

This much hotter back surface temperature reduces the force the feeder can deliver to the melt pool. However, in this deposition mode, the surface tension of the melt pool supplies the adhesion force between the two pieces (as highlighted by the necking behavior seen in Fig 8b).Increasing the feed rate to f=0.75 mm/s results in the opposite effect, decreasing the time for heat to diffuse to the back surface, resulting in a larger temperature difference.

Application SpaceUsing the parameter space defined above for f = 0.5 mm/s, some example pieces were created to showcase the effects of underfeeding the filament during deposition. As shown in Fig 11 below, altering the f/v ratio from 1 (left) to 1/4 (right) shows a significant decrease in layer height, for example variable radius vase structure, from 1.25 mm to 0.45 mm. Transitioning from straight-line depositions to creating complex vase shapes is non-trivial due to additional failure modes and challenges with identifying the correct layer height for the depositions, requiring further work to address these challenges.

While the f/v ratio is not close to the 1/16 target, for this example, we have decreased the equivalent filament radius by half and the layer height by almost 65%.40 W, f=0.5 mm/s, Layer height 1.25 mm55 W, f=0.5 mm/s, Layer height 0.45 mm5 mm5 mm983Figure 11: Comparison of f/v=1 (Left) and f/v=1/4 vases (Right)While a decrease in layer height for artistic vase pieces is useful for demonstration purposes, the technical application space that this technique would be most useful for is in the creation of optical forms, using the decreased layer height to approach the near-net shape more accurately for a lens. As shown in the solid cylindrical pyramid shapes below, there is a significant reduction in the surface variation between the f/v=1 and f/v=1/4 shapes, requiring less laser polishing and secondary post-processing to produce the final desired shape.

This reduction of surface variation is visible in the photographs but most easily identified by viewing the structure profile created from a 3D scan of the pieces (bottom of Fig 12).P=40 W, f=0.5 mm/s5 mm5 mmf/v=1f/v=1/2f/v=1f/v=1/2Figure 13: (Top) Photographs of solid cylindrical pyramid structures at two f/v ratios. (Bottom) Side-view of 3D height map scans of structures (color maps to height of data point)6. ConclusionsThe Digital Glass Forming method has shown the ability to deposit complex 3D glass volumes as fully dense pieces and as sparse freestanding structures. The resolution of this process is limited by the input feedstock diameter, and the high molten viscosity of glass presents challenges for extrusion methods for diameter reduction.

The resolution of this process is limited by the input feedstock diameter, and the high molten viscosity of glass presents challenges for extrusion methods for diameter reduction. Taking inspiration from fiber drawing techniques, the filament is underfed relative to the table scan speed during deposition, but the minimum draw ratio (f/v = 0.2) is limited by a failure mode where the filament lifts off the substrate, stabilizing at some static zError height above the substrate. By shifting the substrate upwards by a ∆z value, this error can be corrected for, allowing for draw ratios as high as f/v=1/16 for straight-line tracks. Altering the ∆z value results in a difference in the temperature profile of the melt pool, maintaining the front surface temperature but decreasing the back surface temperature significantly.

Altering the ∆z value results in a difference in the temperature profile of the melt pool, maintaining the front surface temperature but decreasing the back surface temperature significantly. This decrease in back surface temperature (increase in stiffness) allows the filament to deliver force more effectively from the filament feeder to the melt pool, keeping the filament on the substrate. It is worth noting that the target ratio is not the physical limit of the deposition, only an arbitrary target for this experiment set. Future works on this method will be focused on identifying the physical limits on f/v ratio (e.g., Rayleigh instability or limits in the heat diffusion rate). Additionally, there is a need to estimate the melting pool's loading force as a function of input parameters and explore additional parameter spaces, such as using a different initial feedstock diameter (e.g., 1 mm OD filaments or 125 µm optical fiber).9847.

NTROL PARAMETERS AND MATERIAL SELECTION CRITERIA FOR RAPID PROTOTYPINGSYSTEMSJames W. ComblWilliam R. Priedeman Stratasys, Inc.AbstractSince the introduction of rapid prototyping technology as a tool for time compression and concurrent engineering in the design and manufacturing process, many enhancements and refinements have been made based on the experience ofusers and manufacturers of rapid prototyping equipment. These improvements contribute significantly to faster production of quality output from rapid prototyping systems.There are diverse control and material selection parameters that affect prototypemodels built using the Fused Deposition Modeling (FDM®) process. This paper reviews the role of several of these parameters in the process.

This paper reviews the role of several of these parameters in the process. Data will be presented to help the user choose the appropriate material for specific applications including density, tensile stiffhess, flexural stiffhess, tensile strength, flexural strength, tensile ductility, shock resistance, and hardness.IntroductionWith the commercial Stratasys system now in customer locations for more than twoyears, we have built a substantial base of real life experience with the equipment. The FDM@ process has been an asset to the installed customer base and an acknowledged improvement over previous model building techniques. This experience has prompted design enhancements to better meet the needs of our customers.

This experience has prompted design enhancements to better meet the needs of our customers. As is true for all rapid prototyping manufacturers, we are continually seeking improvements which will deliver more accurate models, of superior surface finish, in increasingly attractive materials, for a better price.Early in 1993, Stratasys released a major enhancement package for the FDM@process which was a direct response to this quest for higher quality models.The intricacies of the control parameters and the interdependency ofthe variables which collectively work to produce models were sorted out in a methodical approach in order to deliver improvements to the existing machine. The FDM@ process allows user control of the envelope temperature, the liquefier temperature, the modeling speeds and the materials to name just a few variables. Each of these variables can alter the resulting model. The appropriate setting ofthese parameters by the operator is key to quality model production.

Each of these variables can alter the resulting model. The appropriate setting ofthese parameters by the operator is key to quality model production. Without proper limits being set, negative results will occur. Additionally, several other features of the system were modified to improve overall performance.86The Stratasys FDM@ process extrudes material via a simple filament drive system. Ifthe capacity ofthe filament drive system is exceeded, the filament can break, bulge, or buckle, causing a plug in the lower filament guide when using certain modeling materials. Other materials may not break, but slipping of the drive wheels may cause improper filament feed (Reference Figure 4a).

Other materials may not break, but slipping of the drive wheels may cause improper filament feed (Reference Figure 4a). To address these issues, modifications of the drive system were developed to increase the overall drive capacity and extensive testing was performed to determine the operating limits to ensure reliable operation.Additional testing was performed to determine the temperature set points for theliquefier and modeling environment for each material. This information led to the re design of the FDM@ process liquefier and cabinet. A heater box package and improved seal system were added to the cabinet to improve the uniformity of the modeling air temperature. A longer and more powerful liquefier was added to increase the volumetric flow rate and improve the temperature consistency of the delivered material.

A longer and more powerful liquefier was added to increase the volumetric flow rate and improve the temperature consistency of the delivered material. Additionally, this liquefier was made to be easily exchanged when changing materials to eliminate cleaning and material "build-up" within the liquefier which contributed to plugging.An improved hold-down method for the modeling foundation was developed toallow the build of larger, thicker models without foundation warping.A firmware solution was implemented to eliminate the "oozing" that occurred at theFDM® tip after shut-off of the material flow. Oozing is the overflow of material that produces small irregularities and loss of detail in the resulting model.Experimental ResultsFigure 1 is a diagram showing the relative increase in traction to drive the filamentthat was achieved with the implementation of the 1/2" elastomeric wheels.

The figure shows two curves: the original 1" steel wheels; the upper curve shows the same for the 1/2" elastomeric wheels. As seen in the figure, the available traction force of the 1" steel wheels is significantly lower than the 1/2" elastomeric wheels. The 1" wheels are more sensitive to changes in filament diameter; Le., normal variations in the filament diameter would produce large variations in the available traction leading to slipping at high material flow rates. The 1/2" wheels, due to their rubber-like behavior, are less sensitive to filament diameter changes and produce more traction in the feed mechanism. This gives the FDM@ filament drive system a higher flow rate capacity.the lower curve is a plot of traction force versus filament diameter forIn Figure 2, the pressure flow relationship for various tip sizes and temperatures are shown. The relationship shown is characteristic ofeach of the modeling materials offered by Stratasys.

The relationship shown is characteristic ofeach of the modeling materials offered by Stratasys. Curves representing liquefier pressure (p)versus volumetric flow rate (V) are depicted. The liquefier pressure is created by the drive traction force acting on the filament divided by the filament's cross-sectional area. For a given tip diameter (di) and liquefier temperature(Ti), the relationship between pressure and flow is roughly linear. As87tip diameter is decreased the pressure required to produce a specific flow rate increases dramatically. As temperature decreases, the pressure required increases due to the increased viscosity of the materiaL In the Stratasys filament drive system the liquefier pressure attained is limited by.: a) the Jorceat which the filament drive slips (Fs); b) the compressive strength ofthe filament (sc); and c) the stress at whichthe filament buckles (sb). The force levels for each ofthese Umits differin.magnitude·and relative order for various materials.

The force levels for each ofthese Umits differin.magnitude·and relative order for various materials. Therefore, the.· systetnis bounded in pressure by the lowest· of these values per material. The system is further bounded by the available liquefier heat exchanger capacity. Ifoperatedbeyond·this maximum flow rate (Vmax), the material delivered will not attain the desired set point temperature.To produce a model, tip. diameter,. process temperature, road width (w), z-slice thickness (z), and speed (s) are selected. The volumetric flow rate is the road width times the z-slice thickness times the speed ofthe head (V=w*z*s). In order to not exceed the operating bounds of the system, the· speed must be selected so that pressure and flow rate remain within the operating limits. VI and V2 represent the maximum allowable flow rates for the upper two curves in Figure 2.The operating parameters discussed above are hard barriers to the modeling process.

VI and V2 represent the maximum allowable flow rates for the upper two curves in Figure 2.The operating parameters discussed above are hard barriers to the modeling process. The system must be operated within these limits to ensure reliable plug-free, slip free operation.Figure 3 is a diagram representing the required set points for liquefier temperatureand air temperature to achieve good models. In general, there are upper and lower liquefier and air temperature limits for each material. Exceeding these limits do not necessarily mean the model will fail but poor surface quality or low part strength may result. Typically, rippling ofthe model surface is caused by the air temperature being set too high and, to a lesser degree, by the liquefier temperature being set too high. Conversely, low modeling air temperatures result in poor bonding strength between the layers for some materials and actual delamination of the model in extreme cases.

Conversely, low modeling air temperatures result in poor bonding strength between the layers for some materials and actual delamination of the model in extreme cases. Low liquefier temperatures result in low limits for material flow rates due to the high viscosity of the material and also poor bonding. Therefore, experiments were conducted for each material to define the guidelines to achieve the optimum balance between surface finish.strength andSystem EnhancementsAs a result of these experiments several design changes were made to the StratasysFDM® process to improve its performance. These changes were delivered to all customers during the first quarter of 1993. This enhancement package consisted of the following design modifications:1. Addition of Seals and Fan Heater Boxes and A New Cabinet Door Design to Improve Air Temperature Uniformity.

This enhancement package consisted of the following design modifications:1. Addition of Seals and Fan Heater Boxes and A New Cabinet Door Design to Improve Air Temperature Uniformity. The combination ofthese three items improves88uniformity of the air temperature within the FDM@ process during the modeling process. This improved air temperature uniformity eliminates cold spots within the environment that could cause poor bonding or delamination ofthe model. Additionally, the existing auxiliary heat circuit used to ramp the cabinet up to temperature now operates automatically.2. Longer, More Powerful Liquefier to Improve Material Delivery and Set Point there is less3. Smaller, Elastomeric Wheels and Larger Filament Diameter to Increase the Buckling Strength and Available Drive Traction. The buckling strength of the filament is a function of its diameter (d) and its compressed length (I).

The buckling strength of the filament is a function of its diameter (d) and its compressed length (I). (Reference Figures 4a and 4b.) An increase of the filament diameter and a reduction of the compressed length increases significantly the filament's resistance to buckling. The decrease in the compressed length was achieved by the use of smaller wheels and the elimination of the lower filament guide. The previous 1" steel wheels were unable to conform to varying filament diameters. Smaller, 1/2" elastomeric wheels are better able to conform to the filament thereby reducing stress concentration and increasing drive traction due to their higher coefficient of friction.4.Improved Hold-Down Tray to Prevent Warping. The Stratasys FDM@ process deposits material on a removable foam foundation. Previous methods to retain this foam base were limited in their ability to prevent the foam from warping during the construction of large, thick parts. The new design.

Previous methods to retain this foam base were limited in their ability to prevent the foam from warping during the construction of large, thick parts. The new design. rigidly holds the foam in an aluminum tray by the use of steel spears. The tray is easily removed from the machine to allow model removal and replacement during pauses in the modeling process. This feature gives the operator the flexibility to perform interim operations on the model not previously possible.5. Enhanced Firmware to Eliminate "Ooze." A roll-back feature was incorporated into the firmware design which rolls back the filament drive wheels at the end of each curve. This feature eliminates the deposition of excess material at the tip, thereby improving the model·quality.89Material SelectionFour different materials are currently available for use with the FDM~ process:1) machinable wax; 2) investment.casting.wax; 3) P200, a polyolefin; and4) P300, a polyamide.

Material selection for a particular model is dependent, in part, upon the end use ofthe model, part design, part size,. and material properties.Models created on rapid prototyping systems typically are used for concept modelsfor design verification and marketing presentations, prototypes for form, fit, and function testing, or patterns for mold making and investment casting. P200 and P300 are most frequently used for concept models and· prototypes while machinable wax and investment casting waxes are used for pattern creation.Part design and size will determine material selection in the building of a part.Good part design reduces the amount of stress in the part and leads to better model quality; Le., less warpage and delamination. In the case of designs that require supports the lamination strength of materials will vary and affect ease of for the build process, support removal.

In the case of designs that require supports the lamination strength of materials will vary and affect ease of for the build process, support removal. Higher strength materials are required where thin wall sections are involved and greater lamination strengths are required for large parts due to the inherent shrinkage factors of the individual materials.Material properties of interest to model builders include tensile strength, flexuralstrength, tensile modulus, flexural modulus, notched impact, unnotched impact, elongation, and hardness.PropertieslMaterialMachinable WaxP300 (polyamide)P200 (Polyolefin)Tensile Strength (psi)1,1141,7651,324Flexural Strength (psi)1,2932,1131,537Tensile Modulus (psi)70,00080,00090,000Flexural Modulus (psi)50,00060,00090,000Notched Impact (ft*lb/in)0.720.240.17Unnotched Impact (ft*lb/in)12.91.461.37Elongation (%)6.653.484.68Hardness (Shore D)407058Table 1.

Material Specifications (based on ASTM tests)90selection of the appropriate material for a model must consider all of theabove factors.We are continually researching new materials with improved properties and modeling characteristics. New materials currently under investigation include powdered ceramics, powdered metals, elastomers, and water-soluble materials.SummaryThe nature and properties of each model are affected by a multitude of modelingparameters. The recent design enhancements to the FDM® process better define and control these modeling conditions and relationships. We will continue to incorporate enhancements and materials into the FDM® system as our knowledge base grows.©Stratasys, Inc.91a>eou... c: o U t\1.=0 III Q) ' - ::::::s C/) C/) a> '- 0- '- Q) ;: Q) ::::::s C'" :......%" Elastomeric Wheels1" Steel WheelsFilament DiameterFllIur.1. Drive traction vs. filament dia.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCONTROL SYSTEM FRAMEWORK FOR USING G-CODE-BASED 3D PRINTING PATHS ON A MULTI-DEGREE OF FREEDOM ROBOTIC ARMAndrzej Nycz*, Mark W. Noakes*, Christopher J. Masuo*, and Lonnie J. Love*Oak Ridge National Laboratory, Manufacturing Demonstration Facility, 2370 Cherahala Blvd, Knoxville, TN 37932AbstractThis paper describes a control system framework using G-Code-based 3D printing paths ona serial link robot manipulator with multiple degrees of freedom. Usually, G-Code is created by a software application, commonly referred to as a slicer, meant for gantry systems. However, G- Code does not address the kinematic complexity nor take advantage of the flexibility available in serial link robot manipulators.

However, G- Code does not address the kinematic complexity nor take advantage of the flexibility available in serial link robot manipulators. This paper provides an overview of the additive manufacturing process and G-Code, types of additive manufacturing deposition movements, common terminology used, the roles of parsers and translators, step-by-step instructions on how to implement this control system, and results and findings from this research. The presented framework can be used for a number of additive manufacturing methods, hybrid solutions, or applications not directly related to additive manufacturing. The implementation was successfully tested on a manipulator with seven degrees of freedom that successfully performed hundreds of hours of large-scale wire arc metal deposition.BackgroundAdditive manufacturing (AM) or three-dimensional (3D) printing is a production process relying on 100% digital and automatic path formulation [1].

The process usually starts with a 3D part model created in a computer aided design (CAD) program. Then, the model is saved in a standard tessellation format or Stereo Lithography (STL) file and processed through a 2.5D path generator. The path generation is commonly called slicing, and the software used to generate these paths is typically called a slicer. Slicing divides a part into vertical layers and beads, which makes going from layer to layer the only vertical movement. A bead is the extruded and then deposited material that forms lines and layers of a part as it is printed. There is usually a start and stop process between each bead as well as between each layer. The output of this process is a G- Code file describing the motion of the printing head and the necessary process settings.G-Code [2] is a computer numerical control (CNC) machine motion programming language. Its roots go back to early 1950s machine tool control and was meant for simple motion control.

Its roots go back to early 1950s machine tool control and was meant for simple motion control. Therefore, it does not take advantage of advanced features available in modern programming languages. Among many of its drawbacks, it lacks looping and variable paradigms. In addition, most of the physical controllers are not designed to execute closed loop control. For most 3D printing processes, the G-Code file is the final path description, and there is no process feedback involved.In AM, the main generator of G-Code is the slicer. The majority of available slicers [3]produce a geometry in Cartesian space that is compatible with gantry systems. However, the1622drawback of using a gantry system in AM is that it limits movements. G-Code generated by slicers gives no consideration for the complexity of kinematics required to control serial link manipulators.

G-Code generated by slicers gives no consideration for the complexity of kinematics required to control serial link manipulators. This makes G-Code-based control of AM manipulators naturally incompatible.A solution to a gantry system’s limited movements is a robotic arm carrying the AMprinting head because it offers increased complexity of motion. However, if the process requires feedback or advanced controls, the G-Code path and process description are inadequate, and an additional translator or post-processing agent is necessary. Additional complexity may occur if the AM process is not omnidirectional. This happens when the printing head cannot be considered a geometric point but is a two or more-dimensional object requiring a specific orientation along the printing path. The majority of fused deposition modeling processes are omnidirectional (Fig.

The majority of fused deposition modeling processes are omnidirectional (Fig. 1), whereas most laser-based and some arc-based wire fed systems are directional and require advanced motion path handling or reduction in the path complexity (Fig. 2).Fig 1. Omnidirectional BAAM-type (Big Area Additive Manufacturing) polymer printing head.IntroductionThis paper describes a framework for translating and executing a G-Code generated path using a seven degree-of-freedom (DOF) Mitsubishi manipulator and advanced tungsten inert gas (TIG)-based wire fed metal arc deposition process as an example.Control of print head orientation is required for proper execution of the build. This paper presents a complete sequence of implementation steps from G-Code reading through translation and execution. The human machine interface (HMI) and a data logger used for post-process analysis are also presented.

The human machine interface (HMI) and a data logger used for post-process analysis are also presented. The principal structure of the described system can be used for any AM process, hybrid additive-subtractive methods, and non-AM applications.1623Fig. 2 Directional GTAW (TIG) deposition head with right side forward feeding wire.There have been a few other implementations of G-Code-based execution for serial linkrobot manipulators; however, none of them are similar to the work presented in this paper. Dritsas and Goulthorpe used computer-aided manufacturing (CAM) software to use a Kuka six DOF robot manipulator to execute milling operations [4].They found the use of G-Code to be quite tedious though it initially appeared to be a simple translation. Human intervention was necessary via an interactive simulation program to modify geometry and machine paths to compensate for the information that G-Code generation lacked.

Human intervention was necessary via an interactive simulation program to modify geometry and machine paths to compensate for the information that G-Code generation lacked. End-effecter orientation and manipulator-specific problems, such as wrapping the arm cables around the arm, were just two of many challenges.Surdilovic et al. [5] similarly used CAM software to generate G-Code to execute millingoperations using a Comau robot manipulator. Like Dritsas and Goulthorpe, they used a simulation program to sort out the problems found in the G-Code that did not adequately describe motion necessary for the manipulator. The simulation software was used to fix collisions, bad manipulability, and joint singularities. The G-Code generation also did not adequately address motion profiles and rapid changes in end-effector orientation.Slavkovic et al. used a five DOF robot manipulator to execute modified G-Code for milling operations [6].

used a five DOF robot manipulator to execute modified G-Code for milling operations [6]. The authors also found that post-processing of the G-Code was necessary to support operation on a robot manipulator. The authors chose to execute an off-line compensation by correcting the G-Code by predicting static tool tip displacements calculated based on a robot compliance model. In summary, it appears that all prior work in the use of G-Code for tool path execution on serial link robot manipulators requires some version of post-processing of the generated G-Code to address the unique needs of each manipulator and the process that is being executed. The work presented in this paper discusses a more general scheme of G-Code post-processing to support both the manipulator and the tooling (AM or machining) processes task that make up operation.1624General System DesignA standard slicer produces toolpath commands controlling the motion of the print head and/or the base table.

The base table is a platform onto which the material is deposited and can move in one or more directions in some systems. To make the algorithm more generic for presentation here, the motion of the table is not considered. However, it can be extended without changing the basic structure. Deposition in layer by layer mode is considered here.Types of Deposition MovesDeposition MovesA deposition move is a move of the printing head from the deposition start point to theend point while continuously printing (depositing material). Traditionally, it is a horizontal path during which the printing head creates a bead (Fig. 3).Fig 3. Deposition moves/beads (green) and travel moves (purple) in a layer of sliced impeller part.Travel MovesA travel move is a type of motion that the printing head takes to move from the end of thejust executed bead to the beginning of the next bead without printing. Consecutive beads typically don’t have adjacent starts and stops.

Consecutive beads typically don’t have adjacent starts and stops. Therefore, they require this additional, nonprinting motion. It is usually comprised of a sequence of up, transverse, and down motions (Fig. 3).Unwind MovesAn unwinding move is necessary when a non-omnidirectional deposition head or a robot manipulator has external cables routed down the length of the arm to prevent hardware damage.Maintenance MovesA maintenance move can be any move required by the printing process such as headcleaning, conditioning, and filament or wire cutting etc.Utility MovesA utility move can be necessary for complex processes and can include scanning,preheating, surface conditioning, or machining.1625System Conditions, States, and AssumptionsThe system shown in Fig. 4, relies on the capability of the programming environment torun many threads at the same time and to provide proper inter-process communication tools such as queues, notifiers, and shared memory [7].

The implementation code for this project is written in LabView and relies heavily on action engines [8], a native LabView feature. Action engines combine the idea of shared memory, semaphores, object programming, and subroutines into one elegant, powerful, and easy to use concept. The presented algorithm can be implemented using different approaches; however, action engines provide a powerful tool for inter-process communication and synchronization.Fig. 4 System diagram.The usual slicer output is a text file (shown in Fig. 5) with a combination of deposition and travel moves organized in layers. Each line represents a single atomic command (Fig. 3 and Fig 5). This text file is the input to the toolpath translator. Depending on the slicer and G-Code generator, the layer’s end beads are defined by key words (G1, G0, etc.) either as part of G- Code language or as custom defined labels. For example, this could be extrusion on/off commands.

For example, this could be extrusion on/off commands. Unless slicer settings are specifically defined for a process from the start, its output might contain redundant information that is often included just to improve human readability.Fig. 5 G-Code example, which is slicer output.1626A good practice before moving the arm to an automatically generated path is to have it parked at a known position and configuration such that the next destination is achievable. This means there are no kinematic singularities or obstacles. For the sake of generality, this location is called Work Home.Unless specified otherwise, the system might require a geometric or process calibrationupdate.

For the sake of generality, this location is called Work Home.Unless specified otherwise, the system might require a geometric or process calibrationupdate. This can be either an automatic or manual process and should be done prior to the start of the deposition process.In the case of large-scale deposition, without going into the specifics of the process, theslicer will not be able to easily produce all the steps and parameters necessary to execute a single node. This is caused by the process’s dynamic nature, arm structure, or closed loop requirement. The alternative is allowing a large number of parts to fail.Terminology and DefinitionsTask (Job)A task or job is the complete set of actions that must occur to build an object/part.NodeA node is a motion destination point described by X, Y, and Z coordinates. The process usually changes at node points in a form of new tool orientation, speed, or deposition.

The process usually changes at node points in a form of new tool orientation, speed, or deposition. A current execution node is a node to which the execution takes the end effector. A previous node is the last node that the robotic arm reached and executed.Node PropertyThe purpose of the node property is to describe specific actions taking place between theprevious and current nodes.ProcessA process is a description of the deposition method. It may dictate a specific armorientation, hardware use, or software interaction.System CalibrationDepending on the type of task and hardware used, the system might have different endeffector or process calibrations. In this work, it was the location and orientation of the print table.LayerA layer is a physical layer of print. On the code level, it is a set of nodes with allcorresponding properties necessary to build a physical layer.TimersTimers are important from a maintenance, testing, and development point of view.

It isimportant to know the amount of time it takes for layers, beads, or non-deposition movements to be completed. It also directly affects the economy of the process [9]. Hence, three basic timers are proposed here: task timer, layer timer, and deposition pause timer. It is also important to note that support code must be concise so that it does not impact execution time.1627StepA step is an atomic execution set consisting of one node, its properties, and all necessaryinstructions to accomplish it. The system adds additional motion or procedures based on an analysis of the step and the arm location as needed.ParserThe parser’s main role is to read the slicer text output and recognize the basic structure elements such as layers and beads.

The parser ignores any commands that are not related to the specific robot manipulator or printing process.TranslatorThe translator receives the data from the parser and packs it into a task consisting oflayers and nodes with assigned properties. In addition, each node carries information of its current layer, the total number of nodes for that layer, and the total number of layers for that task. As necessary, the translator converts the slicer coordinate frame to the manipulator coordinate frame. Applying base height of the manipulator is a good example of such an action.

As necessary, the translator converts the slicer coordinate frame to the manipulator coordinate frame. Applying base height of the manipulator is a good example of such an action. The Parser and Translator could be designed as one entity, but for elegance of design and modularity, it is preferable to split them into two modules.Simulation ModuleAn optional module can be used to run a full kinematics simulation to check if the motionis within the arm’s workspace and to ensure that singularities and link constraints are avoided.Task Execution ModuleThe Task Execution Module (TEM) interprets the newly created structure node by node,applies user interface settings, monitors the process, provides user feedback, executes control loop strategies, and provides input to the data logger.Data Acquisition Module (DAQ)To execute G-Code commands for the given process, it is often important to perform dataacquisition on the robot and process parameters.

For example, data acquisition can be used for closed loop control. This process should be implemented in such a way that it does not interfere with the primary control system while offering adequate acquisition rates.Data LoggerA data logger is a software agent that collects all chosen system settings in a text file witha time stamp. It is used for part certification and build analysis.Graphical User Interface (GUI)The graphical user interface is a critical component of the system that must be addressedwhen creating a control system for advanced AM. GUI enables the user to import geometry, to set proper parameters, and to notify the user of build progress throughout the build. In a case of research and development projects, it is a valuable tool in diagnosing, understanding, and solving problems.

In a case of research and development projects, it is a valuable tool in diagnosing, understanding, and solving problems. Like the data acquisition module, it should be implemented in a separate thread so that the main control systems are not affected by user interface delays.1628System Structure DetailsParser - The parser data flow is as follows (Fig. 6):Fig. 6 Parser flow diagram.1. Read/load G-Code file. 2. Find the start and the end of the useful portion of the file. Not everything in the header and the end of the file is useful for robotic arm motion.3. Identify the first and last line of the layer. 4. Beginning at the start line of the layer, identify all the code lines important for the process using their G-Code command descriptors.5. Identify the move/process or property for each filtered line. 6. Repeat the process until all layers and lines of G-Code are analyzed. 7.

Identify the move/process or property for each filtered line. 6. Repeat the process until all layers and lines of G-Code are analyzed. 7. Insert the information into newly defined structure containing: task name, lines of useful G- Code, and number of layers. The final form for the parser data is presented in Fig. 71629Fig. 7 Parser processed G-Code.Translator - The translator has the following flow steps (Fig. 8): 1. Read the current system calibration. 2. For every layera. Identify each node X, Y, and Z values. b. Apply part calibration. c. Apply default arm orientation. It is usually an orientation with the last joint verticaland pointing down toward the print tabled. If the process is directional, calculate the azimuth angle for the direction.Fig. 8 Translator flow diagram.16303.

If the process is directional, calculate the azimuth angle for the direction.Fig. 8 Translator flow diagram.16303. Create the final task structure in a form following: task name, layers (node vectors: (XYZ), orientation matrix vectors, direction angle vectors, process vector, base height, and total number of nodes in the layers), total layer number, current layer, and current node. This form is presented in Fig. 9. The current node and current layer fields are used during the execution process to keep track of the progress and inform the operator via GUI.Fig. 9 Translator data structure and transformed data.TEM Structure- This module takes one step at a time from the task and executes it in a following manner (Fig. 10): 1. Start new data records (only the first time) with the data logger and data acquisition module. 2. Start task timer. Other timers are handled on an as-needed basis. The most standard timers are the pause timer and layer timer.

2. Start task timer. Other timers are handled on an as-needed basis. The most standard timers are the pause timer and layer timer. Their time data is stored in the data logger and used in the GUI as well as for process analysis.3. Check for Step completion. This portion of the algorithm runs in a logical loop. It means the execution comes back to this point at every step cycle. A new task starts with this flag set to “YES.”4. Check for pause or stop conditions. If the system has pause capability, this portion of the algorithm is used to stop both the motion and the deposition and allow for part inspection while keeping all the system states in the memory. It also allows for clean reentry back into deposition.5. Load the printing speeds and deposition rates from the GUI. 6. Load a Step from the Task. It pulls a single node from the current layer along with its orientation matrix azimuth angle process type and base height.7. Update the GUI, data logger, and data acquisition.

It pulls a single node from the current layer along with its orientation matrix azimuth angle process type and base height.7. Update the GUI, data logger, and data acquisition. All the user-available data gets updated at this step. Data logging and data acquisition receive their start triggers.1631Fig. 10 Task execution module diagram.8. Step execution (Fig. 11). Step execution starts with the application of a variable height. Variable height is optional and advantageous if the process requires manually settable height correction. Next, the system checks for cable unwinding and node properties to determine if immediate actions are needed before the start. If the previous node was in deposition mode and unwinding is necessary, the first action will be to stop deposition. The next step after unwinding is print head (end-effector) orientation. In this part and the position block, the inverse kinematics for the specific manipulator is performed.

The next step after unwinding is print head (end-effector) orientation. In this part and the position block, the inverse kinematics for the specific manipulator is performed. The final part of the sequence is the position command and deposition start sequence. Depending on the process, this can be a series of process-specific steps. If closed loop, real-time control is required for the process or any geometry command over-write, this is the place in the algorithm where it should be applied.9. If this was the last Step, the system moves to final task completion. 10. Next, the algorithm goes back to point 3 and continues until the Step is completed.163211. After all the steps are completed, the system performs a graceful shut down. This includes parking the arm in the Work Home position. If not possible, the arm needs to lift up and clear the created part.

This includes parking the arm in the Work Home position. If not possible, the arm needs to lift up and clear the created part. In addition, all the timers need to be reset, data logging files can now be closed, and the arm should be set to normal settings.Fig. 11 Step flow diagram.System Implementation and ResultsThe presented algorithm was implemented and used to build a wire-arc deposition system. It consisted of a seven DOF Mitsubishi PA10 arm, Miller Dynasty 350 TIG welder, CK Worldwide WF-5 wire feeder, and pyrometers. A TIG torch was mounted on the end effector and used a deposition head. This setup was used to do wire deposition testing to build simple geometric shapes including single and multi-bead walls that were up to 1 ft tall, square tubes, hexagon tubes, and cubes. These geometries are the standard building blocks of any directed energy deposition system. The goal in printing these geometries was to test the feasibly of using1633this system. Eventually, a 1m.

The goal in printing these geometries was to test the feasibly of using1633this system. Eventually, a 1m. long replica of a Willys Jeep bumper (Fig. 12) was printed using this system. The Willys Jeep bumper was chosen to see if it was possible to print something as large as three to four ft. Further mechanical testing of the bumper showed that the isotropic mechanical properties and strength equal a fully solid block of material. With the use of high strength wire, the properties can be as good as or better than original legacy parts.Fig. 12 Willys Jeep with a 3D printed bumper.The practical software implementation was done using LabView (Fig. 13) with before mentioned hardware. The arm-specific control commands were included in the manufacturer supplied DLL (dynamic linked library). A LabView wrapper was created to call these functions and take advantage of the inverse kinematic algorithms.Fig.

A LabView wrapper was created to call these functions and take advantage of the inverse kinematic algorithms.Fig. 13 GUI LabView interface with the algorithm implementation.The GUI interface was updated at a maximum of 4Hz, whereas the TEM had 3ms forceddelays between the cycles with a maximum theoretical rate of 333Hz. The remaining threads were run on demand. Since the system did not require time resolution in micro second range, the implementation was done on a Windows based system with two Xeon CPUs. The low-level commands were executed through Mitsubishi-supplied hardware controllers. Therefore, to take full advantage of real-time capability and closed loop control, a high-level system with 500Hz or better with a hardware clock would be preferable.

Therefore, to take full advantage of real-time capability and closed loop control, a high-level system with 500Hz or better with a hardware clock would be preferable. TEM can be implemented in both software and real-time based systems, but a real-time system is preferred for arc interpolation or complex moves and good process timing.1634ConclusionThis paper presents a system non-specific step-by-step procedure to use slicer-generatedG-Code toolpaths to control robotic manipulators. The presented framework can be used for several additive manufacturing (AM) methods, hybrid solutions, or applications not directly related to AM. The implementation was successfully tested on a seven DOF robotic manipulator that successfully performed hundreds of hours of large-scale wire arc metal deposition.AcknowledgmentThis material is based upon work supported by the U.S.

NTROLLED LOCAL PROPERTIES IN THE SAME PART WITH SINTAFLEX A NEW ELASTOMER POWDER MATERIAL FOR THE SLS PROCESSGideon N. Levy (2) Paul Boehler(1), Raffaele Martinoni(1), Ralf Schindel(2), Peter Schleiss(2) (1) The Valspar Corporation AG, www.valspar.com SWITZERLAND, (2) Institute for Rapid Product Development www.fhsg.ch/rpd University of Applied Sciences St. Gallen, SWITZERLAND Email gideon.levy@fhsg.chAbstract: A new powder material for the SLS (Selective Laser Sintering) process was recently released. The material is a result of fruitful research programs involving industry and university. The well known and widely used DuraFom™ (PA12) and theCastForm™ (PS) SLS-materials were developed by the same team. In the search for new powder materials many properties of the candidate polymer, e.g. the pulverization, the laser absorption and sintering parameters have to be tuned carefully. Previous Elastomer options materials were poor in strength, detailing, and long-term use.

the pulverization, the laser absorption and sintering parameters have to be tuned carefully. Previous Elastomer options materials were poor in strength, detailing, and long-term use. The new product overcomes most of the known deficits. It open completely new practices in many branches like: automotive, house appliances, office equipment, foot ware, medical, and many more. The Sintaflex has a Shore hardness variability 45-75 A and Elongation up to around 300%. The attainable yield strength range is 1.3 - 4.2 MPa. The resolution on the SLS part is up to 0.6 mm. It is positioned in good agreement compared with other commonly used injection plastics. Furthermore, the appeal of all SFF process beside geometry and complexity is thought in varying locally the mechanical properties. Some published patents make suggestions in this direction. The new material, due to the particular properties range in function of the sintering parameter, allows first time to realize this wish.

Some published patents make suggestions in this direction. The new material, due to the particular properties range in function of the sintering parameter, allows first time to realize this wish. The generated part has controlled variable local properties; a new and unique opportunity opens for the SLS process. The paper describes the basic material properties. Further the main sintering parameters are describes and indications on machine settings are given. RP (Rapid Prototyping) applications and the recent practical experience are illustrated. Distinctive examples of local variable properties in the same part and given limits are shown. Some conclusions are stated.Keywords: Rapid Prototyping, Selective Laser Sintering, Elastomer1 The new SLS powder material targeted user needs. We wanted to find an elastomer covering the traditional uses.

We wanted to find an elastomer covering the traditional uses. The typically used final elastic part materials are numerous; starting with natural rubber (NR), butadiene- caotchouc (BR) Nitric caotchouc (NBR) and others, a widely used material in injection molding are the Ethylene-propylene rubbers and elastomers (also called EPDM and EPM). Previous attempt to mimic flexible rubber-like materials was also considered [2]. The previous proposed solution failed due to unacceptable final part properties disintegrating and processing difficulties [10]. Nevertheless user evidenced the need and benefits, the problem remained latent [1]. This research targeted to provide alternatives. As reference properties Neoprem, EPDM, andIntroduction197natural rubber in respect to elongation at break and shore hardness, both the most significant properties in thermoplastic elsatomer material selection, were considered. The final positioning of the new Sintaflex is compared in Fig. 1.

The final positioning of the new Sintaflex is compared in Fig. 1. The appeal of all Layer Manufacturing (LM) processes beside speed, geometry and complexity is thought also in local mechanical properties variation. This feature is also termed gradual properties change. Some published patents make suggestions in this direction. The new material, due to the particular properties range in function of the sintering parameters, allows first time to implement this aspiration. The generated parts may have controlled variable local properties. A new future oriented opportunity for the SLS process.Figure 1. Sintaflex in relation to commonly used elastomers2 Signposts in the Sintaflex research2.1 Basic material It needs a long evaluation and experimental work to validate and formulate a suitable SLS powder material [5]. After fulfilling the SLS process requirement the final parts have to be characterized.

After fulfilling the SLS process requirement the final parts have to be characterized. [3] We find always a radical weakening in the mechanical initial properties, a prediction is not possible, and the deterioration is significant. We have selected finely as basic material Polyester with Shore D 40 [Table 1] & (Fig.2).Table 1: Polyester initial material propertiesTensile stress at break27 MPaElongation at break600%Tensile Impact strength145 KJ/m2E-Modulus172 MPa (at -40°C), 55 MPa (at 23°C), 28 MPa (at 100°CDensity1.1Chemical resistanceResistance to swell in oils, aliphatic and aromatics, hydrocarbons198Figure 2: Initial bulk material shore hardness and the alteration after selective laser sintering2.2 The SLS Powder The bulk basic polymer has to be milled and formulated adding required additives to enable the use on the SLS system. The characteristics of the powder and the grain size distribution have a significant influence on the final properties.

The characteristics of the powder and the grain size distribution have a significant influence on the final properties. Further more, to enable tightness an infiltration seal fluid was developed and is regularly used. Parts in natural state as coming of the SLS platform can be used as well as infiltrated parts (Figures 3, 4). The infiltration is available in several colors.PowderNot infiltrated sintered Sintaflex partsInfiltrated Sintaflex partsFigure 3. SEM pictures of the Sintaflex SLS powder/parts in various stagesFigure 4. The infiltration procedureValspar chose 3D Systems [13] to commercialize the Sintaflex and Sintafluid products under 3D's DuraForm® brand.

The infiltration procedureValspar chose 3D Systems [13] to commercialize the Sintaflex and Sintafluid products under 3D's DuraForm® brand. The products will be known as:DuraForm® Flex plastic LS Material, and DuraForm® FlexSeal (Black, Blue, Yellow, Red, and Neutral) infiltration fluids1992.3 Typical optimized Laser Sintering Parameters The typical sintering parameter for a DTM 2500 plus are given in Table 2.Table 2: SLS typical processing parametersDuring intensive investigations we found out that the most significant parameters influencing the final mechanical part properties are the Laser Power and the Scan Space (SP).3 Mechanical properties SLS sintered Sintaflex partsThe development process is wide-ranging and involves numerous iterations and modifications of the final product. We present here the final optimized properties. Further the influence of the major single parameter, the laser power is given in detail.

We present here the final optimized properties. Further the influence of the major single parameter, the laser power is given in detail. This particular characteristic opens up the new opportunities.The experiments were carried out on a 3D systems Sinter-station 2500+ . In a second step prior to commercialization the optimized parameters sets for all actual Sinter-station platforms were elaborated. A set of typical parameters are shown in Table 2. for orientation only.It was found that the SP scan space has a moderate influence on the achieved mechanical properties mainly the strength (see § 5.1). The most significant properties elongation at break and shore A hardness were found to be well correlated to the laser beam power (LP) during hatching.In order to render tightness the infiltration following process is used. We report on the infiltration later however from here on we have to consider natural parts as coming of the Sinter-station and Infiltrated part post processed.

We report on the infiltration later however from here on we have to consider natural parts as coming of the Sinter-station and Infiltrated part post processed. The diagrams will show the two corresponding states.200Figure 5. Hardness Shore A vs. Elongation [%] for infiltrated and not infiltrated partsIn Figure 5 we can see the shore hardness in function of the elongation with the laser beam power as parameter, the left diagram relates to not infiltrated and the right one to the infiltrated state. In both case we see a clear correlation. It becomes also apparent that we have a properties range. The Sintaflex has a Shore hardness variability 45-75 A the shore hardness, this was measured with a common durometer [4](see also Table 4).For the hardness measurement, in accordance with DIN EN ISO 868, we use a Hildebrand a manual Durometer Model HD3000 with a Drag pointer [11]. It is very important for repeatability and proper results to calibrate the sinter system used.

It is very important for repeatability and proper results to calibrate the sinter system used. The laser power in operating in an open loop mode, the value on the screen mostly does not correspond to the actual real power values. The time dependent drop in laser power output may have significant uncontrolled influence on the hardness and the final part’s mechanical properties.The tensile tests were done on a Zwick testing machine according the ISO 527-2 standard. The mechanical properties E module [MPa] and the tensile strength at rupture Rb [MPa] are quasi linear with the laser power (Figure 6). The influence of the infiltration is very small. Actually from the practical point of view the differences are negligible.Figure 6. The E [MPa] module, the tensile strength Rb [MPa] at ruptures the Laser power [W] interrelations4 The mechanical properties range In summary, the Sintaflex has Shore hardness variability 45-75 [Shore A] with resultant Elongation up to around 300 [%].

The attainable yield strength range is 1.3 - 4.2 [MPa] and the corresponding E module is between 3 and 11.5 [MPa].201The geometrical resolution on the SLS parts is up to 0.6 mm. It is positioned in good agreement compared with other commonly used injection plastics (see also Figure 1). Figure 7 represents some typical parts realized with Sintaflex; many other projects were successfully implemented. [7].Figure 7. Some representative Sintaflex parts (white parts are not infiltrated)The Laser Power [W] parametric variation as shown in Fig. 8, 9 and as said before, controls the main property the hardness and the resultant Elongation [%], E module [MPa] and Rb [MPa],.

8, 9 and as said before, controls the main property the hardness and the resultant Elongation [%], E module [MPa] and Rb [MPa],. Combined with the following infiltration process the final properties are of practical use and concern.nature (green) & infiltrated807013 W15 W15 W6011 W13 W507 W9 W11 WA e r o h S40305 W7 W9 W205 W100300εBruch[%]20010011.522.533.544.55RBreak[MPa/mm2]Figure 8. Hardness, Elongation and rupture tensile strength range with laser power as202parameternature (green) & infiltrated8015 W7013 W11 W15 W609 W13 W507 W11 WA e r o h S40305 W7 W9 W20105 W0300εBruch[%]20010023456789101112Emod[MPa/mm2]Figure 9.

Hardness, Elongation and E module range with laser power as parameterBeyond a valuable new SLS elastomer material we ended up with the first commercialized material with gradual property changeability in a part. The major variable property is the Shore hardness [A] controlled by the laser power used during the selective laser processing.5 Layer manufacturing consistency With the trend and wish to move further towards RM (Rapid Manufacturing,) attention has to be focused on consistent results. Some of the inconsistency results from the process nature. It is a thermal process causing in-homogeneity; it is a layering procedure creating non-isotropic material properties. In the following relevant points related to our case, specifically: material and system are briefly discussed.5.1 Material relatedThe statistical variances measured are shown in Fig 10. In respect to the Rm [MPa] tensile strength and the E module [MPa] the variances are acceptable values.

It is not a suitable test. The not sintered micro voids generate week points in the part (porosity) and may cause undesired part failures. Hence if one prefers consistent strength and high quality over productivity use 0.15mm scan space, the sintering time becomes longer. On the other hand if one wants fast parts with not so consistent strength you may use 0.3mm with better productivity. We recommend using 0.15 mm at all times. In any case users will have to consider the system related calibration for each case (see 5.2)5.2 System related The main parameter is the laser power; unfortunately the laser power deteriorates over time. We have an open loop system. Furthermore the systems have no real power indication on the204control it is just a scalar number. Hence any machine with the same settings may produce different (also unwanted) part qualities.We developed a walk around, a laser power individual machine calibration set.

Hence any machine with the same settings may produce different (also unwanted) part qualities.We developed a walk around, a laser power individual machine calibration set. It consists of standard parts to be produced measured and calibrated by a software program. This calibration enables also to predict the mechanical part properties. It has to be repeated occasionally.A seconded issue is the temperatures settings. We are dealing with a thermal process but have no real absolute temperature indications. Each individual machine has its one temperature values setting. Soft ware temperature deduction methods were implemented.The third missing possibility in standard equipment is the gradual continuous power change in a cross section. A walk around is the subdividing of the cross section into single attached: STL files.

A walk around is the subdividing of the cross section into single attached: STL files. This is only an approximation used for the application examples.It is strongly recommended to remediate these lacunas in future systems and future work in favor of consistent rapid manufacturing applications. [6]6 Gradual properties application example The practical future utilization of gradual properties change in LM produced parts is an innovative challenge. Several researchers are continuously trying to develop patient individual prosthesis using geometrical design features as main local body contact optimization parameter. In reference [8] e.g. a “Double-Wall, Transtibial Prosthetic Socket Fabricated Using Selective Laser Sintering” is proposed. The SLS process is often indicated [9]. Figure 11 described another option. The commercial foot RE (Reverse Engineering) system serves for the precise geometries capture.Figure 11.

The SLS process is often indicated [9]. Figure 11 described another option. The commercial foot RE (Reverse Engineering) system serves for the precise geometries capture.Figure 11. PED-CAD Germany http://www.pedcad.de/beginn.php3This 3-D foot sole measuring device is a step forward. It does not measure only the foot shape precisely but it is: simultaneously during the measuring process, modeling the foot in an ideal shape [12]. The 3-D foot modeler, measures to meet medical and orthopedic technical standards.205In all similar cases we can capture in addition the pressure distribution and local stress distribution. This information can be introduced in the design for finding the best suitable local material attributes e.g. the local hardness. A complementary attribute to the geometrical features normally used. Some research projects in this direction are on the way.6.1 A simple flat sole In order to demonstrate its feasibility a sole was designed, divided in 5 pressure zones.

Some research projects in this direction are on the way.6.1 A simple flat sole In order to demonstrate its feasibility a sole was designed, divided in 5 pressure zones. To each zone a specific shore A value was attributed. The soles were laser sintered on a DTM 2500+ in Sitaflex with the proper hardness related laser power parameters. The infiltrated soles were measured and validated. The results are presented Table 4 and graphically in Fig. 12.5432AB1CDETable 4. Flat sole with locally attributed different hardness and result measurementsZone A B C D E1 73 70 65 59 522 74 70 68 56 513 75 71 69 58 534 74 70 67 55 495 73 70 69 56 52medium 73.80 70.20 67.60 56.80 51.40Sigma 0.75 0.40 1.50 1.47 1.36deviation 1.01% 0.57% 2.21% 2.59% 2.64%80.073.870.267.660.056.851.4A e r o h S40.020.00.0ABZoneCDEFigure 12.

The shore A hardness of a sole in the zones A-E206This new material feature can consequently be applied for other mass customization or product features individualization in various branches.7 Conclusions A new SLS elastomer powder martial was presented. The material enables to vary locally the shore hardens in a certain range. The system limitations and some requirements for safe consistent Layer Manufacturing were stated. An example demonstrating these abilities was shown. Some equipment related modifications are recommended. Supplementary work has to be done on the new materials level. Further system features (HW and SW) and design for layer manufacturing are compulsory.

Supplementary work has to be done on the new materials level. Further system features (HW and SW) and design for layer manufacturing are compulsory. Never the less the unique LM characteristic, gradual properties changes, have future potential.8 References[1] Anonym Somos 201 Helps Reebok Stay On Course, DSM application note, June, 2000 [2] Clausen et al., laser sintering thermoplastic powder US 6.110.411 Patent [3] Childs T:H: Tontowi; 2001; Selective laser sintering of a crystalline and a glass-filled crystalline polymer: experiments and simulations; Proceedings of the Institution of Mechanical Engineering, Journal of Engineering Manufacture; Vol.

ntrolled multi-scale turbulence through the use of Laser Sintered Sierpinski PyramidsLiu, Y., Beck, S., Nicolleau, F. and Majewski, C.E.*, Department of Mechanical Engineering, The University of Sheffield, Sheffield, UKAbstractThe research presented here is the result of a new collaboration between the Centre for Advanced Additive Manufacturing (AdAM) and the Thermofluids group at The University of Sheffield, regarding the use of fractal geometries for the control and influence of fluid flow. It is believed that the use of multiscale objects can be used to introduce many different orders of turbulence into a flow.

It is believed that the use of multiscale objects can be used to introduce many different orders of turbulence into a flow. However, whilst substantial simulations have been carried out in this area, the complexity of the physical geometries means that to date these have not been validated via physical testing.In this work, varying orders of Sierpinski pyramids were produced using Laser Sintered PA2200 and analysed in a wind tunnel with regards to their effects on air flow through the structures. As predicted by theoretical analyses, the coarsest pyramids induced large vortices into the air-stream, whereas the more complex orders induced vortices at a number of different scales, rapidly developing into a standard turbulent flow. Further investigations are planned to isolate the effects of the smaller-scale turbulence in this situation.AimThis paper presents preliminary research investigating fluid flow through 3D fractal geometries.

Previous, theoretical, research makes predictions regarding the effects of different order fractals, but inability to efficiently produce physical models has prevented substantial experimental work. This ongoing research utilises the geometric freedom of Laser Sintering to provide experimental validation of these theoretical studies.Characteristics of turbulent flowFully-developed turbulent flow comprises a range of vortices of varying sizes, similar tothose shown in Figure 1, all of which have an impact on energy transfer.Figure 1 - Example of varied size vortices for turbulence1Whenever a disruption occurs within a flow-stream (for example a physical object such as a valve, or simply a bend in a pipe), this flow structure is likely to change, requiring a certain distance/time before the flow returns to its fully-developed state.

For example, a solid object placed within the flow would be expected to lead to the generation of mainly large vortices, which would gradually break down into the normal range of sizes.In many cases, minimisation of the distance before flow returns to its fully-developed state is required, with the theoretical optimum being the case where the object appears ‘invisible’ to the flow. Whilst difficult to achieve in practice, theoretically this would mean that the flow immediately following the obstruction is indistinguishable from the flow entering it.In other cases, it is more important to control the flow characteristics of a fluid. For example, the accuracy of standard orifice plate flow measurement devices can be sensitive to fluctuations in flow (e.g. asymmetry, flow not being fully-developed), requiring the use of a separate flow conditioner2.

asymmetry, flow not being fully-developed), requiring the use of a separate flow conditioner2. This involves the deliberate placement of an additional object within the flow, upstream of the measurement device, in order to ensure the flow contains the required characteristics.Fractal structuresWhilst the exact origins of fractal-related concepts are debatable, the term ‘fractal’ itself was first proposed by Bernoit Mandelbrot3. Although there are various strict mathematical definitions, at its most simplistic a fractal can be described as a repeating pattern that is self-similar across different length scales.

Although there are various strict mathematical definitions, at its most simplistic a fractal can be described as a repeating pattern that is self-similar across different length scales. Fractal structures are commonly found in nature (for example the human lung, or the branches of a tree), as shown in Figure 2.Figure 2 - Fractals in Nature - human lung, branches of a tree4These structures have also developed applications across a number of sectors, rangingfrom the aesthetic use of mathematical forms5 through to cooling systems6 or computer games7.One of the earliest examples of a fractal geometry, the Koch snowflake8, is shown in Figure3.Figure 3 - Koch Snowflake9The base of this particular geometry is an equilateral triangle. To produce the secondorder fractal, the centre third of each side is replaced with two sides of an additional equilateral triangle.

To produce the secondorder fractal, the centre third of each side is replaced with two sides of an additional equilateral triangle. Whilst only four different orders of the Koch Snowflake are shown in Figure 3, this process can be repeated an infinite number of times, in each case replacing the centre third of each line with a new equilateral triangle.Fractals and turbulenceA number of authors have studied the effects of fractal geometries within a flowstream, whether by experimentation within a wind/water tunnel10,11 or via theoretical modelling12. Of particular relevance here is research into the use of 2D fractal plates as flow conditioners, in combination with an orifice plate flow measurement device, which showed decreased sensitivity to upstream flow disturbances.

Simulation work within the same research also demonstrated that the use of these fractal plates could reduce the length of pipe required to return a flow to its fully- developed state13.Additional work investigated pressure drops across fractal plates, demonstrating theirpotential for use in flowmeters themselves, in place of the standard circular cross-sections14.In order to build on this, and other 2D fractal work, the research presented in this paperinvestigated the effects of several different orders of 3D fractal on downstream air-flow.Experimental workGeometry selectionFor this work, the geometry selected for investigation was the Sierpinski pyramid, anexample of which can be seen in Figure 4.Figure 4 - Sierpinski pyramid15The Sierpinski pyramid is a 3D geometry based around the triangular structures used in theSierpinski sieve16, as shown in Figure 5.Order 4Order 3Order 5Order 2Figure 5 - Sierpinski triangle generationThe structure begins with a solid triangle (order 1).

This can then be divided into four triangles of equal size, and the centre one removed to create the structure seen in the order 2 triangle. This process can then be repeated as many times as required, in each case replacing every solid triangle with this new structure.It is clear that a change in the order of fractal used would necessarily cause a change in the apparent flow area of the geometry. However, for this work the focus was on the effect of different fractal iterations, rather than the effect of variations in flow area. For this reason, the geometries were designed to ensure a constant flow area for each sample.

As it was not possible to maintain a constant overall size and a constant flow area simultaneously, this meant that each pyramid was produced using different overall dimensions, as shown in Table 1.Order/n2 3 4 5Base of total(cm)6.5 7.6 8.9 10.4Base of unit(cm)3.2 1.9 1.1 0.6Height of total(cm)5.6 6.6 7.7 9.0Height of unit(cm)2.8 1.6 1.0 0.6Total volume (cm3) 49 49 49 49Table 1 - Size of Sierpinski pyramidsPart productionThe pyramid geometries were produced in PA2200 (50:50 virgin:used powder) on an EOSFormiga P100 system, using the parameters shown in Table 2.Parameter Laser Power (contour) Scan Speed (contour) Laser Power (hatching) Scan Speed (hatching) Scan spacing Layer thicknessValue 16.0 W 1500 mm/s 21.0 W 2500 mm/s 0.15 mm 0.10 mmTable 2 - Laser Sintering parametersTestingEach sample was mounted within a wind tunnel, and fluid flow achieved by drawing air through the fan shown on the right in Figure 6.

Smoke was generated and incorporated into the air flow, to provide a visible indication of the flow, and a laser/camera system was included to allow easier visualisation of flow and recording of video footage.Figure 6 - Experimental rig set-upResultsThe purpose of this work was to investigate the flow characteristics immediately following the object. Figure 7 shows examples of camera images for each order of pyramid. However, as it is somewhat difficult to observe detail from these still images, Figure 8 provides a schematic of the observed effects.Order 3Order 4Order 2Order 5Figure 7 - Still images of flow through each geometryOrder 3Order 4Order 2Fluid flowFluid flowOrder 5Figure 8 - Schematic of observed effectsWhen observing the Order 2 geometry, it can be seen that the flow immediately following the object is comprised almost entirely of large vortices, of a size comparable with the flow areas (holes) within the pyramid.

As the fractal order of the pyramid was increased, so too the range of vortex sizes increased, again of comparable sizes to the flow areas. In general, the higher order fractals produced downstream flow closer to a fully-developed state than the lower order fractals.Whilst further, in-depth research is planned to investigate this area in more detail, these results demonstrate the potential for the use of this type of 3D fractal to control and or optimise flow characteristics.Future workThe work presented in this paper represents the initial stages of an ongoing investigation into the use of 3D fractals to influence and control fluid flow. Further, more in-depth research is now required in order to establish the capabilities and limitations of these geometries for flow control applications.

Further, more in-depth research is now required in order to establish the capabilities and limitations of these geometries for flow control applications. The planned next stages are:Investigations of different types of fractal geometriesWhilst Sierpinski pyramids were selected for this work, there are a wide variety of other designs which may show even greater potential for control or optimisation of fluid flow. Initial stages will focus on a selection of different 3D fractals, in order to assess their effects on the characteristics of downstream flow. These results will then be applied in order to condition the flowstream either into a fully-developed turbulent flow, or to deliberately induce specific flow characteristics as required.Investigations of specific scales of turbulenceIn this work, fluid flow has been investigates through several orders of 3D fractal.

Whilst the initial focus was on the size, and range of sizes, of vortices within the flow, further information can also be gained. Measurement and comparison of downstream velocity fluctuations in flow through fractals of different orders will allow the determination of information related to each specific turbulence scale.Testing of Kolmagorov LimitThe Kolmagorov Limit is a theoretical size limit below which no additional turbulence scales will be produced, and standard turbulence models break down. However, despite much discussion, this limit has yet to be confirmed. Again, by testing and comparing increasing orders of fractals, it may be possible to experimentally prove or disprove this theory by physical determination of the point at which no further scales of turbulence are observed.As the research focuses on higher fractal orders, increasingly small feature sizes will be required.

NTROLLED-DEPTH LASER CUTTING OF ALUMINUM SHEET FOR LAMINATEDOBJECT MANUFACTURINGGene Zak and Matthew ShiuCentre for Automotive Materials and Manufacturing Department of Mechanical Engineering, Queen's University, Kingston, ON, Canada K7L 3N6AbstractReplacing paper with metal in laminated object manufacturing would bring improvements in terms of part's longevity, and its mechanical and thermal properties. In fabricating laminated parts, a challenging step is blind cutting of the metal sheet. Challenges to overcome are (1) maintaining consistent depth of cut, (2) achieving good surface quality of the cut groove walls, and (3) minimizing the formation of recast. Results are presented of an experimental investigation into controlled-depth laser cutting of aluminum sheet. Thin (0.12-mm) aluminum sheet specimens were cut with a 10-W Nd:YLF laser while varying scan speed and laser power. To accurately observe the cut profiles, specimens were mounted in resin and sectioned.

To accurately observe the cut profiles, specimens were mounted in resin and sectioned. Special handling procedure was developed to handle thin sheet material while avoiding damage. Relationship between cut profile and process parameters was established and shown to conform to established theoretical models.IntroductionUsing laminations (i.e., material in sheet form) for layered manufacturing (LM) avoids phase transformations with attendant shrinkage (e.g., in powder sintering) and allows easy process scale-up. Number of researchers have explored this approach to fabrication of functional parts. Adams et al. [1] produced working tools from steel laminations 1-3 mm thick, but had resort to a finishing milling step to achieve acceptable surface quality. Himmer et al. [2] proposed to fabricate injection moulds by (1) joining flux-coated 2.5-mm-gauge aluminum through heat treatment near the flux melting temperature and (2) finish milling.

Himmer et al. [2] proposed to fabricate injection moulds by (1) joining flux-coated 2.5-mm-gauge aluminum through heat treatment near the flux melting temperature and (2) finish milling. Walczyk and Dolar [3] present results of tests on several epoxy and alumina-based adhesives used to bond aluminum laminations. All of these papers adopt the “cut-stack-bond” approach (Fig. 1a), meaning the process sequence is (1) cut out the desired part’s profile at each layer’s height; (2) assemble (stack) the profiles, and (3) bond them (mechanically with fasteners, adhesively, or by welding). The main disadvantage of the “cut-stack-bond” approach is the complexity of handling the cut-out profiles between the cutting and stacking steps, requiring indexing, transporting, and alignment. This complexity makes automation of the process highly impractical.An alternative approach exists which is more amenable to automation. It is characterized by a “stack- bond-cut” sequence (Fig. 1b).

It is characterized by a “stack- bond-cut” sequence (Fig. 1b). This method has been applied in a commercial LM process, Laminated Object Manufacturing (LOM), but the sheet materials have been limited to paper or polymers [4]. Replacing soft materials with metals is a challenge, however, due to difficulties with “blind” cutting of metals (i.e., cutting only to the depth of the sheet’s gauge).120This paper presents results of experimental investigation into controlled-depth blind laser cutting of aluminum sheet. Before presenting these results, background information on laser-machining of metals is provided.CUTSTACK(a)BONDSTACKBOND (b)Figure 1. Two alternative sequences for producing laminated objects: (a) cut-stack-bond and (b) stack- bond-cut.Laser Machining of MetalsLaser machining of metals is accomplished by achieving a sufficiently high power density (W/cm2) at the focal point to melt or evaporate the material.

The molten material is likely to resolidify, and thus, if most material in the groove is molten, it can produce poor-quality cut, making vaporization the preferred mode.Through cutting vs. blind cutting: In through cutting, the molten and vaporized material can escape through the cut slot; on the other hand, in blind cutting, this material has to be evacuated through the groove’s top opening. Pressurized gas is often used to assist cutting. With through cutting, the gas helps the debris to escape through the cut slot. The role of gas assist in blind cutting is not as significant (and potentially detrimental, pushing the material back into the groove). When cutting metal laminations, the situation more closely resembles blind cutting, since there is no escape route for the molten material and gases through the bottom of the cut.

When cutting metal laminations, the situation more closely resembles blind cutting, since there is no escape route for the molten material and gases through the bottom of the cut. However, there is a difference as well since the groove is being cut not in a bulk material but through a thin sheet with a layer of adhesive belowLaser and optical system considerations: To achieve the high power density required for cutting aluminum (106 - 109 W/cm2) [5], we need to deliver the laser energy over a short time interval (on the order of tenth of nanoseconds) and into a very small area (diameter of 5-20 m m). Most suitable laser beam mode for metal machining is the fundamental (or Gaussian) mode TEM00 because it can be focused to a smallest spot and has the greatest depth of focus compared to other modes. The focal spot diameter is theoretically given by:dl4= D pfwhere l is laser light wavelength, f is the focal length of the focusing lens, and D is the diameter of the incident beam.

The focal spot diameter is theoretically given by:dl4= D pfwhere l is laser light wavelength, f is the focal length of the focusing lens, and D is the diameter of the incident beam. Depth of focus is another important factor to be aware of. It is defined as the distance within which the power density is greater than 90% of its peak value at the focal point. It can be calculated by:Z8 l= pf D22121CUT(1)(2)If Z were too small, even slight surface irregularities and misalignment of specimen surface and X-Y translation plane may cause significant variations in power density. On the other hand, if Z were to be made larger by increasing f or decreasing D, it would also increase the focal spot diameter (Eq. 1), and therefore decrease power density.

On the other hand, if Z were to be made larger by increasing f or decreasing D, it would also increase the focal spot diameter (Eq. 1), and therefore decrease power density. Thus, given the initial beam diameter and laser wavelength, the appropriate choice of beam expansion ratio (affecting D) and focal length involves a trade-off between minimizing the focal spot diameter (d) and maximizing the depth of focus (Z).Laser grooving theoretical models: While much research has been conducted in the area of through cutting of metals [6, 7], the more relevant subject for our work is that of blind cutting (or laser grooving). For example, Modest [8] developed a 3-D conduction model to predict the transient effects in laser machining of ablating/decomposing materials. The model predicts not only the depths but also the shape of a developing groove formed by ablation of material (e.g., aluminum).

The model predicts not only the depths but also the shape of a developing groove formed by ablation of material (e.g., aluminum). Chryssolouris [9] developed a simpler closed-form generalized model to address cutting, drilling, and grooving processes. The model is derived by considering the heat-transfer balance at the constant erosion front produced by the moving laser beam. It predicts the groove depth as:2aP=Ld1 rp 2+TCVd ( s(T o)L)p where a is the material absorptivity (0<a<1), P is the incident laser power, r is the material density, V is the beam scanning speed, d is the focal spot diameter, Cp is the specific and L is the latent heat, and Ts and To are the vaporization and room temperatures, respectively. The relationship implies that the groove depth Ld is directly proportional to the laser power and inversely proportional to the scanning speed.

The relationship implies that the groove depth Ld is directly proportional to the laser power and inversely proportional to the scanning speed. The model assumes that the material is isotropic and its properties are constant; the material changes phase from solid to vapour in one step at a single evaporation temperature; the evaporated/ejected material does not interfere with the beam reaching the erosion front; multiple reflections of the beam within the groove are neglected; and material is removed by vaporization (sublimation cutting).Laser cutting experimentsObjectives: Experiments were conducted with the overall objective of finding suitable processing parameters for controlled-depth cutting of the aluminum sheet material in laminated object manufacturing. The primary controllable process parameters are laser power and scan speed. The primary measured characteristic is the groove depth.

The primary controllable process parameters are laser power and scan speed. The primary measured characteristic is the groove depth. The primary interest in groove depth is due to the need to cut the sheet laminations precisely to the depth of one sheet without significant damage to the underlying stack. A second objective was to determine the effect of changing the material being cut from a (relatively) thick aluminum plate (3.2 mm) to a stack of thin laminated sheets (sheet thickness 0.12 mm). Laser grooving is normally performed on relatively thick material and theoretical predictions would normally also be made for such materials. On the other hand, the material to be cut in our case consists of a sandwich of polymeric adhesive and aluminum sheet. Aluminum is an excellent heat conductor (thermal conductivity 210 W/m C) while a typical adhesive acts as an insulator, with very low conductivity (0.2-0.5 W/m C).

Aluminum is an excellent heat conductor (thermal conductivity 210 W/m C) while a typical adhesive acts as an insulator, with very low conductivity (0.2-0.5 W/m C). Thus, the solid plate represents a “reference” case while the laminated sheet specimens approximates the special material combination to be encountered in the laminated fabrication process.122(3)Specimen Preparation: Two types of specimens were used: solid plates or laminated sheets. The plates were prepared by cutting 2”· 1/8” (50.8 mm· 3.2 mm) aluminum alloy bar stock (ASTM B221 6061 T6511) into 95 mm lengths. These plates were used either directly or as supports for three laminated sheets bonded to them. The latter were cut from a rolled aluminum sheet stock (6061, 0.05” or 0.12 mm gauge) into 90 mm by 20 mm strips.

The latter were cut from a rolled aluminum sheet stock (6061, 0.05” or 0.12 mm gauge) into 90 mm by 20 mm strips. To bond the sheet strips to the support plates, (1) the plates were sanded, (2) plate and sheet surfaces were cleaned by chemical etching, (3) adhesive was applied to the sheets, and (4) using the heat sealer, the strips were bonded one on top of another starting with the support plate. To bond the sheets, a film adhesive was used (Scotch 467MP from 3M consisting of 200MP high-performance acrylic adhesive (50 microns) with polycoated Kraft paper liner).Laser Machining Equipment: The experimental setup consisted of a laser light source, a beam delivery system, and a motion system1. The platform carrying the specimen moved in X-Y plane (with accuracy of – 1 m m) while the optical system remained stationary.

The platform carrying the specimen moved in X-Y plane (with accuracy of – 1 m m) while the optical system remained stationary. The light source was a GSI Lumonics Sigma- 400 diode-pumped Nd:YLF laser operating at a wavelength of 1054 nm (near infrared), capable of up to 12W average power and (Q-switched) pulse frequency up to 20 kHz.The optical path from the laser to the specimen surface consisted of five-time beam expander/collimator, three mirrors, and a 55-mm apochromat triplet lens. Given the light wavelength, initial beam diameter (1.3 mm), beam expansion ratio, and the focal length, the theoretical focal spot diameter is estimated by Eq. (1) to be 11.3 m m. The depth of focus is estimated using Eq. (2) to be only 112 m m, which implies that care must be taken to make sure the specimen surface is parallel with the plane of X-Y motion and that the beam is properly focused.

(2) to be only 112 m m, which implies that care must be taken to make sure the specimen surface is parallel with the plane of X-Y motion and that the beam is properly focused. Assuming actual focal spot diameter to be approximately 15 m m (larger diameter is expected due to optical aberrations of the non-ideal focusing lens), the power density at the specimen’s surface during a single pulse can be estimated to range from 2.3 to 6.8· 109 W/cm2 for laser power outputs of 4W to 12W and pulse frequency of 10 kHz used in the experiments. Therefore, our experimental setup was capable of producing sufficient power density to vaporize the aluminum.Test Procedure: For each specimen, four groups of seven parallel 3.7-cm lines were scanned, with two groups scanned for each power setting (Figure 2). The scan directions in the two groups were reversed with the intention to ascertain absence of any directional effects.

The scan directions in the two groups were reversed with the intention to ascertain absence of any directional effects. (Due to experimental error, for Set 1, both sets of lines for 4 and 7W were drawn in the same direction. Since no significant difference between directions was observed for the two higher powers of Set 1, this error is not expected to affect the results.) Within each group, the scan speed increased from right to left as viewed in the figure, from 1500 to 4500 in 500 mm/min increments. Four power settings were used: 4, 7, 10, and 12W, all at pulse frequency of 10 kHz and pulse duration of 100 ns.Data Collection: After experimenting with other examination techniques (SEM and vertical scanning interferometry, i.e., profilometry), it was found that the only reliable way to obtain the desired information was via metallography.

The difficulties with examining the specimens are caused by the narrow width of the grooves (as little as 15 m m), their fragility, high height to width ratio (up to 8 to 1), and presence of recast and other debris within the groove cut obscuring the groove bottom from non-destructive optical1 Experiments were conducted at the Laser Micromachining Facility, Toronto, Ontario.123examination. The drawbacks of the metallographic examination are that it is destructive and labour- intensive.To make sure that the fragile sheets are completely immobilized, a two-stage mounting process was developed. In the first stage, a 25-mm segment of the support plate with the sheet layers attached was cut away from each side of the specimen (see Figure 2). Each piece was placed in a rectangular rubber mould (85L · 30H mm), and set in epoxy (Araldite GY502).

Each piece was placed in a rectangular rubber mould (85L · 30H mm), and set in epoxy (Araldite GY502). This epoxy block was cut into two pieces, each containing one line group, and, in the second mounting stage, these two pieces were mounted in epoxy within a one-inch diameter cylindrical mould. After polishing to 0.5 m m finish, gray- scale images of sections (256-level, 975· 975 pixels) were captured through an optical microscope and used subsequently to measure the groove cross-section features with a resolution of 0.25 m m/pixel.53W ·51 mm12 mmm m 5 2m m 7 3m m 5 9Section Planesm c 0 5.2Figure 2. Geometry of specimens and line scans (same line pattern was used for all specimens).Results and DiscussionThe observed features of groove cross-sections were groove depth (Ld) and recast area fraction (Ar). The amount of recast within the groove indicates how successful the laser was in severing the sheet.

The amount of recast within the groove indicates how successful the laser was in severing the sheet. Ar is defined as a percent fraction of profile cross-section area occupied by recast. The areas were measured with the aid of Scion Image image-processing software.Set 1: Solid-plate Specimens: Figure 3 shows two examples of groove cross-sections, one with significant recast present and another showing a successful (clean) groove. Both figures show upward molten material flow pattern which developed as the vaporized and molten metal escaped vertically. Recast occurs when molten and vaporized metal resolidifies upon coming in contact with the relatively cool walls of the groove. Under certain process conditions, there is enough resolidification to form a “bridge” across the groove.Figure 4(a) shows the groove depths (Ld1) averaged from two observations. As expected, the groove depth decreases with scan speed and increases with laser power.

As expected, the groove depth decreases with scan speed and increases with laser power. Also shown on the same plot are the results of fitting the model in Eq. (3) to our data. The parameter values used for the model were:124r = 2700 kg/m3, d = 30 m m, Cp = 896 J/kgC, latent heats of fusion 395 kJ/kg and vaporization 10.5 MJ/kg, Ts = 2467(cid:176) C. The model curves were “fitted” by adjusting the absorptivity value (a). The resultant plots are for a = 0.48. Even though this value is higher than the rated absorptivity of aluminum (0.2-0.25), it is well-known that a rises at elevated temperatures [5]. An excellent fit was obtained for all but the lowest power setting. It is possible that the model fails at this setting since it assumes that the material removed from the groove is completely vaporized (i.e., all absorbed incident energy is used to vaporize the groove material). At the lowest power, this assumption may not be valid.(a)(b)Figure 3.

(a) Groove depths (experiment and model) and (b) recast area fraction for solid-plate specimens as a function of scan speed.Set 2: Laminated-sheet Specimens: For laminated-sheet specimens, some combinations of speed and power resulted in successful through cutting of the aluminum sheet (Fig. 5a) while others produced grooves, similar to the solid-plate specimens (Fig. 5b). The trends for the groove depth values (Ld2)1254W7W10W12W5000shown in Figure 6(a) are similar to those observed for the solid-plate specimens except for the upper limit of 120 m m (sheet thickness).Recast area fraction for laminated-sheet specimen (Ar2) displays some of the trends of the solid-plate specimens, but also has some different features (Fig. 6b). In a trend similar to that for the solid-plate specimens, for speeds above 2000 mm/min, the recast area decreases with speed.

6b). In a trend similar to that for the solid-plate specimens, for speeds above 2000 mm/min, the recast area decreases with speed. One important difference, however, is a significant decrease in recast formation at 1500 mm/min for the two highest powers (10 and 12W). These are also the powers and speeds at which a through-cut was observed (see Fig. 5a). It is conjectured that at these higher powers and lower speeds, sufficient energy is delivered through the cut groove to the adhesive below to increase the pressure from vaporized material to the point where the material is removed from the groove before it is able to resolidify.(a)(b)Figure 5.

Examples of groove cross-sections for laminated-sheet specimens: (a) through-cut sheet (12W, 1500 mm/min); (b) recast-free groove (12W, 4500 mm/min).2504 W100.04W2257 W90.07W) s r e t e m o r c i m( h t p e D e v o o r G200175150125100755010W12WSheet Thickness)%( n o i t c a r F a e r At s a c e R80.070.060.050.040.030.020.010W12W10.0250.00 100020003000Scan Speed (mm/min)4000500010001500 20002500 30003500 4000Scan Speed (mm/min)4500 5000(a)(b)Figure 6. Groove depths (a) and recast area fraction (b) for laminated-sheet specimens as a function of scan speed.Comparison of Set 1 and 2 Results: Figure 7 shows the ratio of groove depths for laminated sheet over dy less than one would indicate that, for those for solid plate: equivalent power and speed parameters, shallower grooves were produced in the laminated-sheetL=yL. For example, values ofdd2d1126specimens.

For example, values ofdd2d1126specimens. It is interesting to observe that there is a consistent decline in this ratio with speed for the lowest power (4W), from 0.79 to 0.44, but, for all other (higher) powers, the ratio increases linearly with speed, reaching nearly unity at the highest speeds. Also, at the lowest speeds (1500 and 2000 mm/min), laser power does not appear to affect thedy ratio.1.201.00oi t a R s s e n k c i0.800.60h T0.404W0.207W10W12W0.0010002000300040005000Scan Speed (mm/min)Figure 7. Ratio of groove depths for laminated sheet vs. solid plate (L=ydd2Ld1).ConclusionsFrom laser-cutting experiments on solid aluminum plates and laminated sheets, the relationship between the process parameters (scan speed and laser power) and groove depth was determined and found to agree well with the established theory.

Additionally, a consistent relationship between the groove depths in the solid plate and in the laminated sheets was found. A novel method of metallographic examination of fragile laminated-sheet specimens was developed. Process parameter settings were found which resulted in recast-free through cut of uppermost sheet without significant damage to underlying stack.AcknowledgementsWe gratefully acknowledge the financial support of Centre for Automotive Materials and Manufacturing and Canada Foundation for Innovation. Our appreciation to Laser Micromachining Facility, Toronto, Canada, and to Marc Nantel for his advice. We also wish to express our thanks for assistance with specimen sectioning to Ms. W. Clapp and for microscope imaging to Ms. S. Hinchcliffe, both of Alcan Research Centre Labs.References[1] C. Adams, B. Bryden, D. Wimpenny, “Rapid laminated tooling – making the technology work,” 8th European Conf. Rapid Prototyping and Manufacturing, Nottingham, UK, Jul 1999, pp.

NTROLLING GRAIN EVOLUTION OF IN625 PARTS PRODUCED BYLPBF-AME. Palmer, J. Rosser, M. Ritchie, C. Bromley, S. G. R. Brown and N. P. LaveryFaculty of Science and Engineering, Swansea University, Swansea, SA1 8EN.AbstractLaser Powder Bed Fusion (LPBF) Additive Manufacturing (AM) is rapidly being adopted globally due to its capability of producing complex net-shape parts in a range of alloys with mechanical properties as good as, or better than conventional processes. The alluring possibility is in controlling microstructural features during processing such as grain size, solidification morphology and texture, giving mechanical properties tailored for intended applications. LPBF-AM microstructures are dominated by sizeable columnar growth, which along with even small levels porosity contribute to lower fatigue and creep strength in comparison to wrought. This would limit high temperature applications where IN625 is used in aeroengine exhaust sections.

This would limit high temperature applications where IN625 is used in aeroengine exhaust sections. In contrast, the less dominating equiaxed grains lead to a reduction in crack propagation and improve fatigue performance at the surface. In this work a combination of physical experiments and modelling is used to study the controllability of grain growth and orientation of IN625 made by LPBF-AM.IntroductionOver the years Laser Powder Bed Fusion (LPBF) Additive Manufacture (AM) technologies have received increasing interest across all industries due to the possibility of producing highly functional, near-net shape parts with a minimised amount of post-processing steps.

In particular, the increased degrees of freedom presented by this process has allowed industries, such as the aerospace industry, to produce parts that were previously not possible leading to safer systems with higher efficiencies.Due to the combination of high-strength, fatigue resistance, and anti-corrosion properties [1], IN625 is an excellent candidate for such an industry. IN625 is capable of performing at extreme temperatures for extended periods of time, with resistance to oxidisation and carburisation. This is due to the stiffening effect of niobium and molybdenum on the nickel-chromium matrix, these elements also contribute to the alloy’s resistance and performance in highly corrosive environments.Despite the benefits presented by LPBF-AM, there are key challenges to face. The unique thermal cycle and rapid cooling rates often create complex microstructures and phases that attribute to anisotropic mechanical properties [2], changes in porosity and surface finishes.

The unique thermal cycle and rapid cooling rates often create complex microstructures and phases that attribute to anisotropic mechanical properties [2], changes in porosity and surface finishes. These issues present a new challenge, with the control of grain structures and crystallographic textures in metals being an area of key significance along with the manufacture of tailored parts for specific applications.Previous attempts to control grain structures in LPBF-AM have returned varying levels of success with the focus on changes in scanning strategy, which are often achieved at the expense of build times. The focus of this study is manipulation of laser parameters within the built part via changes in volumetric energy density (V.E.D) and its effect on grain evolution mechanisms. Changes in laser power (W) and speed (m/s) alter the size and shape of the melt pool behind the path of the laser.

Changes in laser power (W) and speed (m/s) alter the size and shape of the melt pool behind the path of the laser. In turn, this effects the temperature gradient (G) along with the liquid/solid interface velocity (R). These parameters were devised by models of heat548transfer and provide good approximation of the columnar to equiaxed transition (CET) of an alloy produced via LPBF-AM [3].Materials and MethodsInconel 625 PowderWithin this work, the IN625 powder used was supplied by Sandvik Osprey with powder size distributions of Dx(10), Dx(50) and Dx(90) values of 25.4, 34.8 and 47.0 μm (measured by particle size analysis), respectively and a composition show in table 1. samples were then built using Renishaw’s AM400 which utilises a 400W ytterbium fibre laser with a 70 μm spot size.

samples were then built using Renishaw’s AM400 which utilises a 400W ytterbium fibre laser with a 70 μm spot size. In addition, the reduced build volume (RBV) auxiliary was also used for rapid testing of various alloys (although, IN625 was the sole focus in this study).Ni Cr MoFeNbAlCSiCo MnTiONPS6221.78.93.83.740.0070.020.02 <0.010.01 <0.010.0090.003<0.003 <0.001Table 1. Chemical composition of IN625 (wt.%)Selective Laser MeltingThere are five controllable parameters that lead to differing volumetric energy densities. Point distance, PD (µm) refers to the distance the laser moves laterally between spots. Hatch spacing, HS (µm) is the parallel distance between two laser lines. Exposure time, ET (µs) refers to the time heat is applied to a singular point. Power, P (W) is the wattage output up to a maximum of 400W in the case of this study. The scanning speed, V (m/s) is calculated by dividing point distance by exposure time.

Power, P (W) is the wattage output up to a maximum of 400W in the case of this study. The scanning speed, V (m/s) is calculated by dividing point distance by exposure time. The oxygen level in the build chamber was reduced to under 0.1% by introducing Argon gas creating an inert atmosphere.The volumetric energy density, V.E.D (J/mm3) can be defined as the volume of energy input to the material per millimetre cubed and is calculated through equation 1 and describes the relationship between these parameters. layer thickness, LT (µm) remained at a constant 60 µm for all builds.𝑉𝐸𝐷 =𝑃 𝑉. 𝐻𝑆. 𝐿𝑇=𝑃. 𝐸𝑇 𝑃𝐷. 𝐻𝑆. 𝐿𝑇(1)For this study, nine density cubes were built using differing parameters as seen in table 2. A tenth configuration was also built using the suggested parameters set out by Renishaw.

𝐻𝑆. 𝐿𝑇(1)For this study, nine density cubes were built using differing parameters as seen in table 2. A tenth configuration was also built using the suggested parameters set out by Renishaw. The substrate under each cube was the optimised configuration named the crucible.Density OptimisationEach cube was analysed by a method based on Archimedes principle; density measurements were taken for all 10 samples using the sigma 700/701 force tensiometer. Water was used as the suspension fluid but was combined with an emulsifier so that breaks in surface tension would not result in inaccurate values for density. Temperature of the air and water were recorded, with initial mass values and change in mass after each test, from this, density in g/cm3 could be calculated along with porosity from the theoretical density.

Density could then be549calculated as a percentage with the normalised density being calculated from the densest sample, assuming it to be 100% dense.For further verification, the cubes were then sectioned and prepared for optical porosity analysis using the Zeiss Observer to view the XY and XZ planes in full view. Scanning electron microscope (SEM) imaging was undertaken with a JEOL JSM-6010PLUS.ImageJ was used to analyse the captured images. This was done by formatting the image and adjusting the threshold so that porosity was the area of interest. Readouts can then be given for average size and area of porosity, working backwards from this, density measurements can be obtained for each sample. The results from these are given in table 3.Table 2.

and Tsai, N. (1983) "Temperature Fields Produced by Traveling Distributed Heat Sources" [4]. Melt pool geometry was calculated using both measured and sourced thermophysical properties (see Table 3), namely, absorption (A), density (ρ), specific heat capacity (Cp), the solidus temperature (TS), the liquidus temperature (TL), the critical temperature (Tc) and the thermal conductivity (k). The solution uses a gaussian heat distribution in dimensionless form to indicate final melt pool shape to accurately predict weld pool shapes for specified alloy systems. This was computed using a MATLAB script and compared against experimental data.A0.46ρ (kg/m3)8440Cp (J/kg.K) 451.67TS (K) 1581.05TL (K) 1653.65Tc (K) 9460k (W/m.K) 9.36Table 3. Thermophysical properties of Inconel 625 (solid state)From these simulations values for the temperature gradient (G) and solidification velocity (R) could be found.

Thermophysical properties of Inconel 625 (solid state)From these simulations values for the temperature gradient (G) and solidification velocity (R) could be found. The temperature gradient is the gradient of temperature across which growth morphology occurs in crystalline metals. Given by:550𝐺 = 𝛥𝑇(2)The solidification velocity is the ratio of the solidification cooling rate within the metalto the temperature gradient given by:𝛿𝑇 𝛿𝑡=𝑇𝑆 − 𝑇𝐿 𝑡𝑆 − 𝑡𝐿(3)where TL and TS are the liquidus and solidus temperatures reached at times tL and tSrespectively.Giving solidification velocity (R) as:𝑅 =1 𝐺𝛿𝑇 𝛿𝑡(4)Layered BuildsFurther to the original 10 builds, a layered sample was produced that consisted of a cube built with a high energy density with regions of lower energy density embedded in eight layers of varying thicknesses. Each layer is an 8 x 8 x X mm where X is varied from 0.1-1 mm.

Each layer is an 8 x 8 x X mm where X is varied from 0.1-1 mm. The high and low energy density layers contained the same parameter sets as A3 and A6 respectively, this is shown in figure 1.Figure 1. layered build showing energy density layers (left) dimension of layers (right, dimensions in mm).Metallographic PreparationEach sample was removed from the build plate using a bandsaw to cut through the supports, a BUEHLER IsoMet 4000 precision saw was used to face the samples, with the density cubes being cut in the XY and XZ directions and the layered build being cut in the XZ direction as this was parallel to the grains being observed. Each sample was mounted via hot compress in conductive bakerlite with a BUEHLER SimpliMet XPS1 mounting system. The samples were then polished with a BUEHLER AutoMet 250 and EcoMet 250 pro grinder- polisher.

The samples were then polished with a BUEHLER AutoMet 250 and EcoMet 250 pro grinder- polisher. The procedure was as follows: 400 to 1200 grit Si grinding pad until plane, then 9, 6 and 1μm water-based diamond polish with polishing pad each for ten minutes. The layered sample was then etched in a mixed acids reagent, 15ml:10ml:10ml - Hydrochloric Acid, Acetic Acid and Nitric Acid. The samples were immersed for 5 to 10 seconds at room temperature before washing off with ethanol and drying. The density cube samples were not etched as analysis of the porosity was the focus in contrast to the microstructure.551Results and discussionDensity OptimisationThe experimental values are shown in table 3. Figure 2 shows experimental density ofboth Archimedes and optical compared with V.E.D.Figure 2 shows a near uniform density within the builds, with A3 providing the greatest density and the Renishaw optimised settings producing near 100% density when normalised with the other builds.

The lower density samples of A6 and A8 typically correlate with a lower volumetric energy density indicating more porosity within the build. There appears to be an anomalous result with the A9 build, due to having one of the lowest V.E.D it was expected to have a greater amount of porosity but appears nearly 100% dense when a density of 96-98% was expected.Figure 2. Graph of volumetric energy density (J/mm3) vs density (%)Table 4. Values of V.E.D and densitySample LabelV.E.D (J/mm3)Density (%)Density Normalised (%)Density from ImageJ (%)A145.92103.4399.73099.784A262.50103.4199.75299.849A379.37103.49100.00099.913A471.43103.1999.53199.802A558.59103.2099.77299.839A636.4699.0495.55794.442A769.44102.9599.40099.153A841.67101.1497.53597.796A937.04103.2099.57199.610AR54.42103.1499.98699.689Figures bellow show the surface of density cube samples A6, A3 and AR.

As expected, there is a direct and visible correlation between V.E.D and porosity, with higher V.E.D leading to less porosity. No microstructure can be seen in these images without etchants. A6 possesses the greatest amount of porosity of the three builds, with large volumes of inclusions created by552escaping gasses during the build process. At low V.E.D, there is often incomplete laser coverage causing high levels of porosity. High density samples such as A3 still contain small amounts of porosity but often with large areas of no visible porosity. Scratches are often seen on the surface; these are caused by the dislodging of un-melted powders within inclusions which are dragged across the surface when polishing.

Scratches are often seen on the surface; these are caused by the dislodging of un-melted powders within inclusions which are dragged across the surface when polishing. These scratches could be removed with thresholding in ImageJ as to not skew the data.For sample A3, the density calculated from ImageJ is largely similar to the tested density but is slightly lower due to the visibility of porosities, whereas the tested density had been normalised to 100% due to being the densest sample. For the Renishaw build, the tested density and density from ImageJ were largely concurrent, with ImageJ slightly overestimating. There is a large disparity with the A6 samples of around 4%. The captured images were in the XY plane parallel to the build plate, due to these porosities being defects from build parameters, the captured images may exaggerate the volume of porosity as there is no gauge of depth in the images.Figure 3.

A6 XY (left) A3 XY (centre) AR XY (right)Electron Microscopy and EBSDFigures 4(a) and (b) show the A6 sample at varying magnifications in both the XY and XY directions. These samples had been etched prior to imaging. Within these images some porosity can be seen and it’s important to see how microstructural changes occur when interacting with porosity and inclusions. Figure 4(a) shows the melt pool profiles and ‘fish mouthing’ caused by the outline of the melt pools, both figures 4(a)and (b) show the cellular dendritic growth of the microstructure within each layer, this growth is common within LPBF- AM alloys [5] and is often referred to as constrained growth. Both coarse and elongated cells can be seen in figure 4(b), often adjacent to each other. Figure 4(b) shows these different microstructures broken up via coloured lines.

Both coarse and elongated cells can be seen in figure 4(b), often adjacent to each other. Figure 4(b) shows these different microstructures broken up via coloured lines. It is postulated that this type of growth is caused by competition between Molybdenum and Niobium as they separate on cooling causing a hexagonal sub structure [6]. Future work may include an EDS map to verify the composition of these grains. The elongated cells show epitaxial growth over the fish mouthing where the coarser cells do not. This is due to the re-melting of several layers during frequent passes of the laser. This has allowed for the grain growth process to restart in each subsequent layer [7].

This is due to the re-melting of several layers during frequent passes of the laser. This has allowed for the grain growth process to restart in each subsequent layer [7]. There is a theory that the coarse cells don’t experience overlapping with various layers, so are not remelted by subsequent laser passes and don’t experience the same energy input in the form of heat, causing a stall of the grain growth process leading to the formation of a coarser microstructure, often referred to as unconstrained growth [8]. It can also be seen how porosity stops the formation of elongated cells whereas the coarser, more hexagonal cells tend to conform around them. This agrees with the previous argument of layer remelting as the escape of gas causing porosity would be several layers deep causing no remelting of the layers and subsequently no grain growth.

This agrees with the previous argument of layer remelting as the escape of gas causing porosity would be several layers deep causing no remelting of the layers and subsequently no grain growth. The irregular inclusion seen in figure 4(b) may have been caused553by a defect in the build such as a line defect due to the parameters used, the lack of laser passes would explain the absence of elongated cells but the presence of some hexagonal cells.Figure 4. SEM Image of A6 XY (a) SEM Image of A6 XY, green lines show porosity, red lines show distinction between coarse and elongated grains, blue lines show the melt pool profile, yellow lines show secondary phases within grains (b).EBSD Layered SampleFigures 5 and 6 contain the images of the etched layered sample containing layers of energy densities corresponding to A3 and A6. The image is characterised by alternating light and dark bands corresponding to differing volumetric energy densities.

The image is characterised by alternating light and dark bands corresponding to differing volumetric energy densities. The lower V.E.D layers are characterised by light bands containing large amounts of porosity. The higher V.E.D is slightly darker with a more uniform surface texture. The greatest presence of observable microstructures occurs at the boundaries between different V.E.D and in the bulk of the thicker layers of greater energy density. Coarser cells are more common within the lower energy density layers but are also seen in the higher energy density layers. As seen in figure 5, the layers are encased within a boundary of the higher energy density, this extends to every face of the sample and can be thought of as a cube within a cube. There also appears to be more microstructural features in the surrounding perimeter giving rise to the elongated cells and the darker blue colour.

There also appears to be more microstructural features in the surrounding perimeter giving rise to the elongated cells and the darker blue colour. This is due to the greater number of laser passes in those specific areas, imparting more heat energy causing elongated cells and grain growth.As the presented images are in the XZ plane, these cells are nucleating upwards, perpendicular to the build plate. In “Fundamentals of solidification” by W. Kurz, Kurz postulates that these elongated cells grow along the direction of the thermal gradient and in a preferred crystallographic orientation, whereas the smaller, coarser like grains could grow between the thermal gradient and their preferred orientation [9].

This concurs with what can be seen in figure 6, elongated cells are most present on the melt pool boundaries and tend to grow perpendicularly into the melt [10] and often transition to more equiaxed coarser grains within the melt pool boundaries, this would be along the thermal gradient of the melt pool. As previously stated, the elongated cells can grow epitaxially across melt pool lines but are stopped almost instantly when reaching a region of differing volumetric energy density, whereas there are a few examples of coarse equiaxed grains crossing over melt pool lines, but they tend to be uncommon.554Figure 5. Optical image of layered sample XZ (left) with accompanying EBSD of layers (right).Figure 6. Optical image of layered sample XZ top centre (left) top left (right).This layered sample was also analysed via EBSD, shown in figure 5, and largely agreed with the optical analysis seen.

Optical image of layered sample XZ top centre (left) top left (right).This layered sample was also analysed via EBSD, shown in figure 5, and largely agreed with the optical analysis seen. Columnar grain growth was more apparent in the regions of high energy density and when undisturbed, climbed through the part in the build direction whereas this growth was inhibited in the lower energy density region where equiaxed grains were formed. In addition, porosity was also seen to affect grain grown in a similar way. Columnar growth is paused below the pores and equiaxed growth begins above it as seen in this SEM image.555Simulation WorkMelt pool dimensions were taken from experimental values and those given computationally within MATLAB from Eagar and Tsai’s solution, across a range of scan speeds and powers (table 4), this data is shown in figure 7.

There were a broader range of depth values for those taken experimentally compared to the values taken from MATLAB, there may be two reasons for this. The plots produced by MATLAB include the entire heat effected zone of the melt, so only values above the solidus temperature should be taken to calculate the melt pool size. It is important to note that the Rosenthal solution used in these simulations does not account for the latent heat of transformation within the set metal parameters along with the nonlinear effects of temperature-dependent properties [11]. This produces temperature estimations vastly greater than the evaporation temperature of IN625 which were not achieved during the practical builds hence everything above the liquidus temperature was largely ignored.

Other issues arise when comparing against captured images, there are large amounts of overlapping of melt pools within the build which is to be expected but means accurate calculation of melt pool depth was more difficult with greater variance. Two value sets are given for the depths calculated via Eagar and Tsai’s solution which relate to the laser absorptivity values used. As this value was unknown, a sensitivity study was conducted which revealed a value of 0.8 gave more accurate depth measurements. A data set using a value of 0.2 has also been included to show the variability between possible parameter values.Table 5.

The determination of these states is based on where values fall around two mathematically derived curves called Hunt criterion lines. These curves are based on statistical probability of grain growth from an initial dendrite. If a value falls below the equiaxed dendrite line at which the volume fraction of equiaxed dendrites is greater than 49% this area corresponds to a fully equiaxed grain morphology. If a value exceeds the columnar dendrite line, at which the volume fraction of equiaxed dendrites is less than 0.66%, the grain morphology is seen as fully columnar. Between these two lines the grain morphology is estimated to be a mixture of columnar and equiaxed.557Using equations 2-4 in conjunction with the data from MATLAB, a CET graph was able to be plotted, seen in figure 8. It is shown that as V.E.D is increased the likelihood of columnar growth becomes more apparent, conversely, application of a lower V.E.D would result in a mixture of columnar and equiaxed.

It is shown that as V.E.D is increased the likelihood of columnar growth becomes more apparent, conversely, application of a lower V.E.D would result in a mixture of columnar and equiaxed. This is largely agreeable with both the optical and EBSD analysis of the layered sample, with energy densities corresponding to A3 (79.37 J/mm3) exhibiting strong columnar growth and energy densities corresponding to A6 (36.46 J/mm3) exhibiting more equiaxed growth. For all power and speed combinations modelled, the thermal gradient saw very little variation compared to the solidification velocity. The thermal gradient was expected to decrease with increasing depth and increasing solidification velocity. It is also shown how solidification rate is highly dependent on process parameters.10000000Increasing V.E.D1000000Columnar100000MixedEquiaxed)m K/10000(G10001001010.0000010.000010.00010.001 R (m/s)0.010.11Figure 8.

CET Plot of thermal gradient and solidification velocity against a range of process parameters.ConclusionIN625 was used in conjunction with LPBF-AM to investigate grain morphology growth mechanisms during applications of varied energy density. The process parameters were first optimised through means of density measurements, first based on Archimedes principle, then optically through ImageJ. This provided two main parameter sets with one high volumetric energy density corresponding to A3 (79.37 J/mm3) and one low volumetric energy density corresponding to A6 (36.46 J/mm3), Renishaw’s optimised parameter set AR was also used in comparison with the optical density measurements.Firstly, scanning electron microscopy was used as a microstructural analysis tool in effort to understand how porosity effects grain growth, this revealed two main sub-cellular cell558300350400structures, coarser and elongated grains.

The presence of these grains was due to the heat energy imparted by the laser which differed based on the volumetric energy density of each build. These structures exhibited different growth mechanisms across melt pool lines and around porosity.Secondly, a layered build consisting of alternating layers of volumetric energy in varying thicknesses was produced. This enabled different grain growth mechanisms to be seen when crossing boundaries of differing volumetric energy density, it was shown optically that columnar grains grew preferentially in higher V.E.D, but their growth was stopped immediately when reaching a boundary of two differing layers. More equiaxed grains grew in both high and low energy densities and exhibited some ability to grow across the boundaries of two layers but this tended to be infrequent.

More equiaxed grains grew in both high and low energy densities and exhibited some ability to grow across the boundaries of two layers but this tended to be infrequent. EBSD was carried out on the layered sample and agreed with what was seen, columnar grain growth was more apparent in the regions of high energy density and when undisturbed, climbed through the part in the build direction whereas this growth was inhibited in the lower energy density region where equiaxed grains were formed.Finally, simulation-based work was carried out within MATLAB to measure melt pool size and shape, but some disparity was observed when comparing with values measured experimentally. The melt pool profiles produced via MATLAB included the entire heat effected zone, not just the liquid melt. Values taken experimentally were often crowded with other melt pool lines making distinction and measurement more difficult.

Values taken experimentally were often crowded with other melt pool lines making distinction and measurement more difficult. Using MATLAB, values for the thermal gradient (G) and the solidification velocity (R) were calculated and plotted on a columnar to equiaxed transition (CET) curve. It was shown that as V.E.D is increased, the likelihood of columnar growth becomes more apparent, conversely, application of a lower V.E.D would result in a mixture of columnar and equiaxed grains, this was in agreeance with what was seen both optically and through EBSD on the layered sample.References[1] V. Shankar, K. Bhanu Sankara Rao, S.L. Mannan, Microstructure and mechanical properties of Inconel 625 superalloy, Journal of Nuclear Materials, Volume 288, Issues 2–3, Pages 222-232. 2001.[2] V.A. Popovich, E.V. Borisov, A.A. Popovich, V.Sh. Sufiiarov, D.V. Masaylo, L.

O. Seabra, M. F. Vieira, Mechanical and microstructural characterisation of bulk Inconel 625 produced by direct laser deposition, Materials Science and Engineering: A, Volume 838. 2022.[11] S. Bontha, The effect of Process Variables On Microstructure in Laser-Deposited Materials M.S., Wright State University. 2002.560.

NTROLLING MELT POOL DIMENSIONS OVER A WIDE RANGE OF MATERIAL DEPOSITION RATES IN ELECTRON BEAM ADDITIVE MANUFACTURINGEmrecan Soylemez and Jack L. Beuth Department of Mechanical Engineering, Carnegie Mellon University Pittsburgh, PA 15213Karen Taminger NASA Langley Research Center Hampton, VA 23681AbstractElectron beam-based additive manufacturing processes are being seriously considered for manufacturing and repair applications in the aerospace industry. To be successful, these processes must work over a wide range of material deposition rates to combine affordability (requiring high deposition rates) with the ability to precisely deposit fine geometries (requiring low deposition rates). Melt pool size and shape are key characteristics to control in these processes. Control of melt pool dimensions will greatly increase the ability to successfully build shapes, and may play an important role in controlling solidification microstructure.

Control of melt pool dimensions will greatly increase the ability to successfully build shapes, and may play an important role in controlling solidification microstructure. In this paper, we present an analytically-guided approach for maintaining melt pool cross sectional area and thermal finite element simulation results are presented over a wide power range (1-5kW) to evaluate the approach. Single bead finite element simulations include the effects of temperature- dependent properties, latent heat, material addition and the distribution of power by a rapidly moving beam. Experiments were carried out on electron beam deposition equipment at NASA Langley Research Center and results show the same trends as those seen in the models.

Experiments were carried out on electron beam deposition equipment at NASA Langley Research Center and results show the same trends as those seen in the models. Ultimately, a map of curves of constant melt pool cross sectional areas and length-to-depth ratios is presented, covering power and velocity ranges over roughly a factor of 5.IntroductionLaser and electron beam-based additive manufacturing processes are being seriously considered for manufacturing and repair applications in various industries. At the low power range, Laser Engineered Net Shaping (LENS®), which has been developed by Sandia National Laboratories and commercialized by Optomec®, has been used to create complex prototypes, tooling, and small-lot production items with the ability to manufacture shapes based on geometry from CAD solid models [1]. Using a similar approach, Electron Beam Manufacturing (EBM) processes are being considered for manufacturing aerospace components.

Using a similar approach, Electron Beam Manufacturing (EBM) processes are being considered for manufacturing aerospace components. Like their laser-based counterparts, in EBM, parts or features are built up layer-by-layer with the beam serving as a moving heat source [2]. Electron beam-based processes offer added advantages over laser-based processes, including more efficient energy transfer to the substrate, transfer efficiencies that are not a function of the reflectivity of the substrate, the ability to deposit a wider range of materials, and the ability to rapidly move the electron beam across the surface or within the melt pool to locally tailor surface temperature fields.An analytical solution related to these processes was derived by Rosenthal [3] and consists of a concentrated heat source moving across the surface of a semi-infinite half space.

In this paper, Rosenthal models fitted to replicate results from finite element simulations are used to help interpret results and guide the selection of numerical simulations to perform. Previous work by the authors and their co-authors on beam-based SFF processes dates back more than a decade. Early work addressed the modeling and measurement of residual stress in metal and polymer deposition SFF processes [4-7]. Control of melt pool size under steady-state conditions for laser- based SFF processes has been considered by using a process map approach to plot melt pool size over the full range of process variables for LENS® [8-11]. The process map approach has been extended to consider melt pool size control under transient conditions and as a function of process size scale [12-15]. In the work on transient melt pool response, numerically determined melt pool response times are used to establish a lower bound on the response times of thermal feedback control systems.

In the work on transient melt pool response, numerically determined melt pool response times are used to establish a lower bound on the response times of thermal feedback control systems. In the work on process size scale, the role of process size on the sensitivity of melt pool size to fluctuations in beam power or velocity are related to process robustness. The issue of residual stress control for laser-based SFF has also been addressed using a process map approach [16, 17]. In that work, a defined thermal gradient behind the melt pool is proposed as a means for predicting changes in final magnitudes of residual stress based on thermal simulation results. The work of [18] applies this approach to the concept of stress reduction by localized part preheating via a dual-beam laser or electron beam system.Where this paper takes a broad process control approach to understanding EBM processes, other studies have focused on complex behavior inside the melt pool.

A mathematical model of laser penetration during laser welding under keyholing conditions is presented in [19]. A 2-D axisymmetric incompressible, viscous transient model was coded to include the effects of surface tension and vaporization in [20]. Detailed mathematical models of melt pool behavior were derived in [21, 22] that include surface tension and natural convection inside the melt pool.Material microstructure plays an important role in EBM process control, due to mechanical property requirements in finished parts. Rosenthal analytical calculations and finite element simulations have been used to observe cooling rates and thermal gradients, and their relation to grain morphology [23]. A study of process variable (beam power and velocity) effects shows that microstructure can vary significantly along the depth of the deposit, and that under high power conditions a transition from columnar to mixed or fully equiaxed structures is possible [24].

It was shown experimentally in [25] that melt pool size and cooling rate significantly depend on the travel velocity and laser power.Sophisticated process models include the effect of material deposition on heat transfer conditions. In [26], 3-D simulations were performed using a dummy material method to model the addition of SS410 within the LENS power range. Following that, 2-D simulations were improved including wire feed effects [27]. This approach is similar to that used in the numerical modeling of this study, but the focus here is on mapping out melt pool dimensions over a broad power range (roughly 1-5kW) of interest in EBM.Experiment-based optimization studies have been performed to optimize melt pool cross section size [28]. In this paper we consider the control of melt pool dimensions of cross sectional area and length via modeling that spans the full range of EBM power and travel speedvalues of interest.

In this paper we consider the control of melt pool dimensions of cross sectional area and length via modeling that spans the full range of EBM power and travel speedvalues of interest. Model predictions were followed by experiments for single bead deposition at NASA Langley Research Center.NomenclatureT baseT Tm k ρ c L Q or P beam power α V d Area melt pool maximum cross section areatemperaturebase plate temperature (and initial substrate temperature)melting temperature thermal conductivity density specific heat melt pool length behind the maximum depthfraction of beam power absorbed by the substrate beam travel speed melt pool maximum depthAnalytically-Based Approach To Find Constant Area For Single Bead DepositionIn this paper we will consider single bead deposition, presenting it as a building-block for the understanding of melt pool dimensional control in depositing complex geometries.

In the EBM processes considered herein, material is added to the molten pool via a wire feed mechanism. Material deposition alters the geometry of the substrate and should be considered for its effect on melt pool size and shape. In general, melt pool sizes will be larger if the geometry of the deposited bead is included, due to a less direct conduction pathway into the substrate.As will be shown in this paper, however, a Rosenthal-based analytical approach which neglects material addition, temperature dependent thermal properties and other physical effects of relevance in EBM processes can still be used as a guide to understand how changes in beam powers and velocities affect melt pool dimensions. Specifically, we use Rosenthal models with thermal properties chosen at a single temperature between Tbase and Tm to allow the matching of a melt pool dimension with that predicted by a single sophisticated thermal finite element simulation.

This “fitted Rosenthal” model is then used to predict the dependence of that melt pool dimension on beam power and velocity. Among other things, this allows the careful choosing of power and velocity combinations to model with the more cumbersome finite element models, which can take days to execute.The goals of modeling single bead deposition are to identify baseline process variables for desired material deposition rates and bead geometries (as a quick reference guide for processing engineers), to identify process variable changes needed to maintain bead geometries as deposition rates are changed over wide ranges, and to greatly increase operating ranges. Achieving the final goal will allow consistent build conditions to be achieved under high material deposition rates for rapid, economical building of parts, and under low deposition rates for use in regions requiring geometric precision.Figure 1 shows fitted Rosenthal analytical results for single bead deposition of Ti-6Al- 4V.

Each curve in the plot corresponds to beam power and velocity combinations yielding a constant melt pool cross sectional area. Values of melt pool cross sectional areas of 0.016, 0.031 and 0.063 in2 are considered. Values for total melt rate, which is sum of molten material deposition rate and the rate of substrate melting, are provided in boxes for each of the 3 curves in 1kW intervals (from 2kW to 5kW). The curves of Fig. 1 were obtained by two fitted Rosenthal models over two power ranges. For beam powers at the substrate (αQ) larger than 3500W, Ti-64 thermal properties at 1523 K were used. For beam powers at or below 3500W, properties at 1200 K were used. Because of this, a slight discontinuity is seen in the shapes of the curves at a power of 3500W.Figure 1. Analytical predictions for power and velocity combinations yielding constant cross sectional areasThe curves of Fig.

Analytical predictions for power and velocity combinations yielding constant cross sectional areasThe curves of Fig. 1 can be used by processing engineers by first specifying a material deposition rate in lb/hr or in3/min (9.5lb/hr is approx. 1 in 3 /min for Ti-64). For a given feed wire diameter, this fixes the wire feed rate. Next, a ratio of molten deposited material vs. substrate material must be specified. In work to date, a ratio of 1:0.30 has been used. This then determines the total rate of melting (for matching to values in the boxes). One finding of this work is that the total rate of melting is largely determined by the beam power. At each power level, the total melt rate values do not vary significantly between the green, red and blue curves (though the differences are greatest at 2kW). The final step in using the plot of Fig. 1 is to specify a melt pool (bead plus melted substrate) cross sectional area.

The final step in using the plot of Fig. 1 is to specify a melt pool (bead plus melted substrate) cross sectional area. This determines what curve to use, and determines the travel speed, V. Another way of thinking about the results of Fig. 1 is that the total rate of material melting is largely determined by the beam power. Beam velocity then determines whether you have a large melt pool moving slowly across the plate or a smaller melt pool moving rapidly across it, but in all cases the total rate of melting remains roughly the same.The procedure outlined above allows use of the results of Fig. 1 to choose beam power, Q, and travel speed, V, for user-specified values of deposition rate, ratio of molten deposited material to substrate material, and melt pool cross sectional area. However, the plot is also a guide to maintaining melt pool area as beam power, beam velocity and material deposition rates are changed.

However, the plot is also a guide to maintaining melt pool area as beam power, beam velocity and material deposition rates are changed. To maintain bead cross sectional area while changing material deposition rate, beam power and velocity are increased or decreased while staying on a given curve. At the same time, wire feed rate is increased or decreased proportionally with beam velocity.There are other ways that the results of Fig. 1 can be used. For instance, given a cross sectional image from a single bead deposition experiment with a length scale, it can be determined what processing curve deposition occurred on, based on bead and remelt cross sectional areas. Also, if a material deposition rate is given, travel velocity can be determined by dividing by the bead cross sectional area. Beam power can then be determined from the plot.It is interesting to note that the curves of Fig. 1 are nearly straight lines.

Beam power can then be determined from the plot.It is interesting to note that the curves of Fig. 1 are nearly straight lines. Thus, relatively simple rules of thumb can be formulated for use by processing engineers. The curves of Fig. 1 are not through the origin, however. This is due to a finite melt pool cross sectional area that exists as the velocity is reduced toward zero. A commonly used rule of thumb for e-beam or laser-based processes is that one can maintain deposition conditions over a range of powers and velocities by keeping the ratio of beam power to beam velocity constant. This is typically referred to as the “P/V = constant” criterion. If this criterion were correct in predicting melt pool areas, the curves of Fig. 1 would be straight lines through the origin. The consequence of the results of Fig. 1 is that a P/V = constant rule of thumb can give very poor results in small-scale processes (e.g.

1 would be straight lines through the origin. The consequence of the results of Fig. 1 is that a P/V = constant rule of thumb can give very poor results in small-scale processes (e.g. in the < 1kW range) but it will become more accurate in understanding melt pool area control for large scale processes. In such cases, the offset of the curves from the origin has less of an effect.Numerical Approach To Predict Constant Melt Pool Area For Single Bead DepositionThree-dimensional finite element simulations have been performed to more precisely model single bead deposition, and to relate to the analytical results presented in Fig. 1. Two kinds of simulations were performed to obtain melt pool dimensions. The first type, designated as an added material simulation, uses the model change function in ABAQUS to directly simulate the addition of material by the wire feeder.

The first type, designated as an added material simulation, uses the model change function in ABAQUS to directly simulate the addition of material by the wire feeder. In this type of simulation, all elements are defined and built initially, but elements in the added material portion of the model are deactivated using the remove function. After that, at each step, elements are added to simulate the addition of new bead material. In the current version of these models, the bead material existing during each load step extends 3 elements ahead of the heat source, which aids in the numerical solution and models heat transfer into some of the wire as it is fed into the melt pool. The second simulation type is designated as a constant bead model. In these models, the entire bead geometry is included in the simulation at all times. This greatly simplifies the modeling, and still captures many of the geometric features of the added bead.

In these models, the entire bead geometry is included in the simulation at all times. This greatly simplifies the modeling, and still captures many of the geometric features of the added bead. However, this allows heat transfer to occur well ahead of the melt pool in the bead material. Thus this type of model will tend to yield, for the same processing conditions, slightly smaller melt pool sizes than the more accurate added material simulations.Consistent with deposition in a vacuum, these 3-D models do not include convective heat transfer on their vertical and top surfaces. The models have an initial temperature of 373 K and a constant temperature of 373 K specified at the base. Eight-noded linear brick elements are used throughout the model.

The models have an initial temperature of 373 K and a constant temperature of 373 K specified at the base. Eight-noded linear brick elements are used throughout the model. A distributed heat flux along the top of the added bead is used instead of concentrated heat flux, more closely modeling oscillatory movement of the electron beam during deposition (which is used in experiments at NASA Langley to reduce the concentration of heat on the surface).Figure 2(a) provides a far-field view of the 3-D mesh used for the symmetric added material and constant bead simulations. To reduce computational time, the mesh is biased toward the middle of the model, where steady-state melt pool sizes are determined. Figure 2(b) shows melt pool thermal contours from an added material simulation.

Figure 2(b) shows melt pool thermal contours from an added material simulation. Figure 2(c) illustrates the geometry of the added bead and melt pool maximum cross section shape, through a cross sectional view (in this case the heat source is moving out of the page).(a) Added material and constant bead 3-D model geometry and mesh (b) Added material melt pool image(c) Maximum cross section area of the melt pool (image is a cross sectional cut view of the geometry shown in (a) and (b))Figure 2. Added material simulation images and temperature distribution near the melt poolNumerical results for constant bead and added material models are given in Table 1 for comparison with Rosenthal analytical results used to generate the plot in Fig. 1. The melt pool cross sectional area values given in Table 1 show that, for high velocity cases (blue line), thereare no differences between constant bead and added material model results.

1. The melt pool cross sectional area values given in Table 1 show that, for high velocity cases (blue line), thereare no differences between constant bead and added material model results. For low velocity cases (green line) noticeable differences are seen. This is because, at high velocities, heat diffusion has no time to distribute heat ahead of the melt pool, so the additional material in the constant bead model does not affect the results. In comparing added material and Rosenthal analytical results, a similar trend is seen. High velocity (blue line) cases show very good correlation between the analytical and numerical models. Low velocity (green line) results show a significant difference, with the analytical results always being smaller than the corresponding numerical results. In these large cross sectional area cases, the geometry of the bead is having an effect on the results. The results of Table 1 thus indicate that the curves plotted in Fig.

In these large cross sectional area cases, the geometry of the bead is having an effect on the results. The results of Table 1 thus indicate that the curves plotted in Fig. 1 are accurate representations of melt pool behavior, except for the green curve results, and particularly for the low power, low velocity cases on that curve.Table 1. Comparisons between Rosenthal analytical results, constant bead model simulation results and added material simulation resultsControl Of Melt Pool Length To Depth RatiosMelt pool cross sectional area is a key geometric quantity to control in EBM. It is related to bead cross sectional area and to melting depth into the substrate. Although the link between melt pool area and depth is not direct for cases of material addition (it depends on the bead geometry), for the Rosenthal model it is direct, with A = π d2/2 (melt pool depth is simply the radius of the semicircular melt pool cross section).

A second key geometric quantity is melt pool length, here defined as the length along the surface from the point corresponding to the maximum melt pool depth to the melt pool trailing edge. Experiments have shown that melt pool length can affect solidified deposited bead shape, consistent with its determining the geometry of the solidification front along the back of the melt pool. Along curves of constant melt pool area plotted in Fig. 1, melt pool lengths can change by a factor of more than 2. These facts have motivated a second study of melt pool length to depth or L/d ratios.Table 2 presents added material numerical results for L/d ratios taken along the three constant melt pool cross sectional area curves of Fig. 1, with the curves corrected to agree with the more accurate added material finite element simulations. Using these results, a second order polynomial equation for L/d has been fitted along each of the 3 curves.

Using these results, a second order polynomial equation for L/d has been fitted along each of the 3 curves. These polynomial fits where then used to determine curves for constant L/d ratios, intersecting the curves for constant melt pool cross sectional areas. The final result of this work is presented in Fig. 3, which gives curves of not only constant melt pool areas, but also curves of constant L/d, all determined from added material finite element simulations.Table 2. L/d ratios obtained from added material simulation resultsαQ (W)Velocity (in/min)L/dGreen LineRed LineBlue Line2000 3000 5000 2000 3000 5000 2000 3000 500010.1 14.1 23.2 17.1 25.8 44.5 32.8 51.5 86.73.16 4.63 6.27 3.38 5.13 8.80 4.66 7.14 11.11As shown in Fig. 3, curves of constant L/d are very much different than curves of constant melt pool cross sectional area, and, in fact, are nearly normal to the middle (red line) curve corresponding to a cross sectional area of 0.031 in2. The plot of Fig.

The plot of Fig. 3 now allows a processing engineer to not only identify beam powers and velocities to yield a desired melt pool cross sectional area, but to also then vary power and velocity to maintain that cross sectional area, or to maintain the current L/d ratio. Results shown in Fig. 3 demonstrate that for a specified material deposition rate (which roughly determines beam power) the high velocity (blue) curve yields the largest L/d values.Figure 3. Lines of constant melted cross sectional areas and L/d ratios as determined by added material finite element simulationsAnalogous predictions for L/d have been obtained using Rosenthal analytical calculations, yielding curves for constant L/d ratios. However, separate fitting of the results is needed for melt pool lengths and melt pool depths. To calculate melt pool length, Ti-64 properties at 992 K were used. For melt pool depth values, properties at 1523 K were used for cases with beam powers larger than 3500W.

To calculate melt pool length, Ti-64 properties at 992 K were used. For melt pool depth values, properties at 1523 K were used for cases with beam powers larger than 3500W. For beam powers at or below 3500W, properties at 1200 K were used (identical to that used for melt pool areas because the two dimensional quantities are linked for a Rosenthal model). Predictions were compared to those from addedmaterial simulations plotted in Fig. 3. The comparison showed similar trends in results, but numerical values were not in good agreement. This was due to differences in melt pool lengths themselves, and due to differences in depth values for the same cross sectional areas (due to geometric differences). It was found that multiplying all Rosenthal-derived L/d values by a factor of 1.2 gave reasonable agreement with the added material simulation results in most cases, but not all.

It was found that multiplying all Rosenthal-derived L/d values by a factor of 1.2 gave reasonable agreement with the added material simulation results in most cases, but not all. Overall, the use of fitted Rosenthal-based analytical formulas to predict melt pool cross sectional areas seems fairly robust. However, trying to use such a model to predict anything more than trends in melt pool lengths is not accurate.Experimental ResultsA collaboration between Carnegie Mellon and NASA Langley Research Center has allowed testing of the predictions provided in Fig. 1. Experiments were performed for single bead deposition using beam power and velocity combinations along each of the 3 curves presented in Fig. 1. Specimens were sectioned at two locations along the bead length in the steady-state portion of the bead. Figure 4 shows a typical cross sectioned bead image.

1. Specimens were sectioned at two locations along the bead length in the steady-state portion of the bead. Figure 4 shows a typical cross sectioned bead image. Note that the boundary of the remelted portion of the substrate (designated by red arrows) is not the boundary of the entire discolored region of the substrate. That boundary designates the boundary of the heat affected zone.Figure 4. Cross section micrograph showing the deposited bead, the remelted substrate boundary (see arrows) and the substrate heat affected zoneExperimental results for melt pool cross sectional areas are presented in Fig. 5. As the figure indicates, melt pool cross sectional areas were relatively constant. The exception to this finding is the lowest power case for the low velocity (green) line that is uppermost on the plot. Although this case is inconsistent with the Rosenthal-based curve for constant cross sectional area provided in Fig.

Although this case is inconsistent with the Rosenthal-based curve for constant cross sectional area provided in Fig. 1, it is very much consistent with the more accurate finite element predictions presented in Table 1. This shows the value of including the geometric details of bead deposition in some cases. The agreement between the absolute numerical values for melt pool cross sectional areas and the experimental values is not very good, with experimental values always being higher (sometimes significantly so) than the predicted values. Investigations are underway to try to pinpoint the source of this disagreement. Regardless, however, the trends in the experiments clearly indicate the ability to control melt pool cross sectional areas over a very wide range of beam powers.Figure 5. Experimental results presenting total (deposited+melted) cross section areas with changing power.

Experimental results presenting total (deposited+melted) cross section areas with changing power. Velocities were selected according to fitted Rosenthal calculations to obtain a constant area.ConclusionsIn this paper, analytical and numerical methods are used to map out combinations of beam powers and beam velocities for controlling melt pool dimensions in electron beam manufacturing (EBM) processes. The melt pool dimensions considered are melt pool cross sectional area and the ratio of melt pool length (L) to maximum depth (d). Results are presented which show specific paths through processing space to yield constant melt pool areas or L/d ratios. Experimental results are presented which show a clear ability to maintain melt pool cross sectional areas over a wide range of practical powers.

Experimental results are presented which show a clear ability to maintain melt pool cross sectional areas over a wide range of practical powers. Modeling results are also presented that demonstrate why a P/V = constant rule of thumb for maintaining melt pool dimensions can yield poor results, particularly for small-scale processes.AcknowledgementsThe authors wish to acknowledge collaborations with Bryant and Raymond Walker of Keystone Synergistic Enterprises, Port St. Lucie, FL for additional experiments and physical insights related to this research. This research was supported by the National Science Foundation under grant CMMI-0700538, the Turkish National Ministry of Education and the Nick G. Vlahakis Graduate Fellowship.References1. Hofmeister, W., Griffith, M., Ensz, M., and Smugeresky, J., “Solidification in Direct Metal Deposition by LENS Processing,” JOM, September 2001, pp. 30-35.2. Taminger, K. M. B.

Hofmeister, W., Griffith, M., Ensz, M., and Smugeresky, J., “Solidification in Direct Metal Deposition by LENS Processing,” JOM, September 2001, pp. 30-35.2. Taminger, K. M. B. and Hafley, R.A., “Characterization of 2219 Aluminum Produced by Electron Beam Freeform Fabrication,” Solid Freeform Fabrication Proceedings, Proc. 2002 Solid Freeform Fabrication Symposium, Austin, August 2002, pp. 482-489.3. Rosenthal, D., “The Theory of Moving Sources of Heat and Its Application to Metal Treatments,” Transactions of ASME, Vol. 68, 1946, pp. 849-866.4. Klingbeil, N.W., Zinn, J.W. and Beuth, J.L., "Measurement of Residual Stresses in Parts Created by Shape Deposition Manufacturing," Solid Freeform Fabrication Proceedings, Proc. 1997 Solid Freeform Fabrication Symposium, Austin, August 1997, pp. 125-132. 5.

1997 Solid Freeform Fabrication Symposium, Austin, August 1997, pp. 125-132. 5. Klingbeil, N.W., Beuth, J.L., Chin, R.K, and Amon, C.H., "Measurement and Modeling of Residual Stress-Induced Warping in Direct Metal Deposition Processes," Solid Freeform Fabrication Proceedings, Proc. 1998 Solid Freeform Fabrication Symposium, Austin, August 1998, pp. 367-374.6. Ong, R., Beuth, J.L. and Weiss, L.E., "Residual Stress Control Issues for Thermal Deposition of Polymers in SFF Processes," Solid Freeform Fabrication Proceedings, Proc. 2000 Solid Freeform Fabrication Symposium, Austin, August 2000, pp. 209-218.7. Klingbeil, N.W., Beuth, J.L., Chin, R.K. and Amon, C.H., "Residual Stress-Induced Warping in Direct Metal Solid Freeform Fabrication," International Journal of Mechanical Sciences, Vol. 44, 2002, pp. 57-77.8. Vasinonta, A., Beuth, J.L. and Griffith, M.L., "Process Maps for Laser Deposition of Thin- Walled Structures," Solid Freeform Fabrication Proceedings, Proc.

44, 2002, pp. 57-77.8. Vasinonta, A., Beuth, J.L. and Griffith, M.L., "Process Maps for Laser Deposition of Thin- Walled Structures," Solid Freeform Fabrication Proceedings, Proc. 1999 Solid Freeform Fabrication Symposium, Austin, August 1999, pp. 383-391.9. Vasinonta, A., Beuth, J.L., and Ong, R., "Melt Pool Size Control in Thin-Walled and Bulky Parts via Process Maps," Solid Freeform Fabrication Proceedings, Proc. 2001 Solid Freeform Fabrication Symposium, Austin, August 2001, pp. 432-440.10. Vasinonta, A., Beuth, J.L. and Griffith, M.L., "A Process Map for Consistent Build Conditions in the Solid Freeform Fabrication of Thin-Walled Structures," Journal of Manufacturing Science and Engineering, Vol. 123, No. 4, 2001, pp. 615-622.11. Beuth, J.L. and Klingbeil, N.W., "The Role of Process Variables in Laser-Based Direct Metal Solid Freeform Fabrication," JOM, September 2001, pp. 36-39.12.

123, No. 4, 2001, pp. 615-622.11. Beuth, J.L. and Klingbeil, N.W., "The Role of Process Variables in Laser-Based Direct Metal Solid Freeform Fabrication," JOM, September 2001, pp. 36-39.12. Aggarangsi, P., Beuth, J.L., and Griffith, M.L., "Melt Pool Size and Stress Control for Laser-Based Deposition Near a Free Edge," Solid Freeform Fabrication Proceedings, Proc. 2003 Solid Freeform Fabrication Symposium, Austin, August 2003, pp. 196-207.13. Birnbaum, A.J., Aggarangsi, P. and Beuth, J.L., "Process Scaling and Transient Melt Pool Size Control in Laser-Based Additive Manufacturing Processes," Solid Freeform Fabrication Proceedings, Proc. 2003 Solid Freeform Fabrication Symposium, Austin, August 2003, pp. 328-339.14. Aggarangsi, P., Beuth, J.L., and Gill, D.D., “Transient Changes in Melt Pool Size in Laser Additive Manufacturing Processes,” Solid Freeform Fabrication Proceedings, Proc. 2004 Solid Freeform Fabrication Symposium, Austin, August 2004, pp. 163-174.15.

2004 Solid Freeform Fabrication Symposium, Austin, August 2004, pp. 163-174.15. Birnbaum, A.J., Beuth, J.L. and Sears, J.W., “Scaling Effects in Laser-Based Additive Manufacturing Processes,” Solid Freeform Fabrication Proceedings, Proc. 2004 Solid Freeform Fabrication Symposium, Austin, August 2004, pp. 151-162.16. Vasinonta, A., Beuth, J.L. and Griffith, M.L., "Process Maps for Controlling Residual Stress and Melt Pool Size in Laser-Based SFF Processes," Solid Freeform Fabrication Proceedings, Proc. 2000 Solid Freeform Fabrication Symposium, Austin, August 2000, pp. 200-208.17. Vasinonta, A., Beuth, J.L., and Griffith, M.L. “Process Maps for Predicting Residual Stress and Melt Pool Size in the Laser-Based Fabrication of Thin-Walled Structures,” ASME Journal of Manufacturing Science and Engineering, Vol. 129, No. 1, 2007, pp.

NTROLLING PHASE COMPOSITION IN SELECTIVE LASER MELTED STAINLESS STEELSThomas L. Starr1, Khalid Rafi2, Brent Stucker2 and Christopher M. Scherzer3 University of Louisville, Louisville, KY USA 1Chemical Engineering, 2Industrial Engineering, 3Rapid Prototyping CenterAbstractCommercially important stainless steels can be austenitic or martensitic and this phase composition fundamentally controls the mechanical properties of the material. With selective laser melting (SLM), 17-4 stainless steel can be produced in either phase depending on powder composition, SLM conditions and post-build heat treatment. This behavior is examined using optical and electron microscopy and high temperature x-ray diffraction in order to better understand the formation of metastable austenite and its transformation to martensite.

Control of phase composition can produce a material with either extremely large strain-to-failure or high yield strength and can provide a method for completely eliminating residual stress.IntroductionStainless steel with a nominal composition 17% Ni, 4% Cr, 4%Cu is a common alloy for commercial, defense and medical applications. With conventional processing (casting and solution heat treating) the 17-4SS alloy is martensitic at room temperature and can be precipitation hardened to provide high yield strength and good elongation. Previous investigations with 17-4SS using SLM have reported various results with respect to as- processed phase composition. Facchini (2010) and Starr (2010) found that SLM produced a metastable austenitic material that cannot be hardened with standard heat treatment but that transforms to martensite during mechanical straining and exhibits exception elongation.

Starr (2011) also reported that heating the as-process material to 788oC and cooling produced martensite which could be hardening conventionally. More recently Murr (2011) found that the as-processed phase composition depends on the gas used during powder production and on the atmosphere used during the SLM process. Better understanding is needed both to ensure reproducible manufacturing of hardened stainless steel parts and, perhaps, to exploit the unique mechanical properties of the metastable austenitic form of this alloy.ExperimentStainless steel powders with nominal 17% Ni, 4% Cr were obtained from two sources:EOS (EOS GmbH, Munich, Germany) and LPW (LPW Technology USA, Maumee, OH). Both powders are formed using the gas atomization process with particle size and morphology suitable for use in the SLM process. The EOS powder was atomized in nitrogen, the “standard” atmosphere for stainless steel powder manufacturing.

The EOS powder was atomized in nitrogen, the “standard” atmosphere for stainless steel powder manufacturing. The LPW powder was a “special order” and was atomized in an argon atmosphere. Composition and particle size for the two powders are shown in Table 1.Test specimens were created using an EOS M270 Dual Mode Direct Metal LaserSintering machine, capable of operating under either nitrogen or argon atmosphere. All439specimens were fabricated using standard EOS- specified exposure parameters with a layer thickness of 20 or 40 micrometers. Exposure parameters for these two layer thicknesses are shown in Table 1.Table 1. SLM exposure parameters Layer (μm)2040Received powder and fabricated specimensPower (W)195195were characterized by a variety of methods. The particle size distributions of the powders were measured using a Microtrac 3500 Particle Analyzer (Microtrac Inc, Montgomeryville, PA).

The particle size distributions of the powders were measured using a Microtrac 3500 Particle Analyzer (Microtrac Inc, Montgomeryville, PA). Carbon and nitrogen content of EOS powders and fabricated parts were measured using combustion analysis (ASTM 1019-08) by IMR Metallurgical Services (Louisville, KY).Speed (mm/s)Spacing (mm)Energy density (J/mm3)1000 8000.10 0.109861Powder morphology and dense specimen microstructure were observed using optical(Olympus MX 5) and scanning electron microscope (FEI Nova Nano SEM). The specimens for optical and SEM were prepared using standard metallographic specimen preparation methods. 10% oxalic acid electrolytic etching and modified Fry’s reagent were used to etch the specimens.Mechanical properties specimens were tested in tension using an Instron 5569A 50kN test machine (Norwood, MA) under displacement control at 1.0 mm/min.

Stress-strain curves were obtained from load cell and strain gage output using Instron’s Bluehill software.The martensite content of fabricated specimens was measured by the magnetic induction method using a Feritscope FMP30 (Fischer Technology, Inc., Windsor, CT). This technique measures the total amount of ferritic phase in austenitic and duplex steel. The technique cannot distinguish between α-ferrite, δ-ferrite, and martensite (α’). For 17-4SS formation of ferrite is not expected and the Feritscope reading is assumed to the percentage martensite. Before each measurement session the instrument is calibrated against steel standards with 2.5, 10.5, 30 and 100% ferrite. Since the SLM materials are stainless steel the magnetic response is different than that of these carbon steel standards.

Since the SLM materials are stainless steel the magnetic response is different than that of these carbon steel standards. A fully martensitic (“condition A”) bar of commercial 17-4PH measured 75% ferrite.X-ray diffraction analysis utilized a Bruker D8 diffractometer system with a dome- covered stage that allows heating of specimens in the range 30o C to 1100o C during data acquisition. For high temperature experiments a borosilicate glass slide was used over the ceramic heating element to prevent the x-ray diffraction pattern of this alumina stage from interfering with specimen pattern. Specimens for XRD analysis were cut parallel and perpendicular to the build plane using a low speed diamond saw and polished in three steps from course to fine culminating in a 0.5 micron diamond polished finish. After acquiring a room temperature pattern, each specimen was heated to a temperature of interest, held at this temperature for one hour and then cooled at 1o C/s to 200o C.

After acquiring a room temperature pattern, each specimen was heated to a temperature of interest, held at this temperature for one hour and then cooled at 1o C/s to 200o C. Additional XRD patterns were obtained at fixed temperature over the range 200o C to room temperature at 20o C increments. The duration of each scan was roughly 11 minutes. The total time per run for each specimen was approximately three hours.440Results and DiscussionThe results below show the effect of the SLM process and the post-build heat treatmenton phase composition, microstructure and mechanical performance of 17-4 stainless steel.As-received PowdersTable 2 shows the oxygen and nitrogen content of both powders. As expected, the EOSpowder, atomized in N2, has substantially more nitrogen than the argon-atomized LPW powder. Figure 1 shows XRD traces of the as-received powders. The EOS powder is mostly austenite (γ), while the LPW powder is mostly martensite (α’).

Figure 1 shows XRD traces of the as-received powders. The EOS powder is mostly austenite (γ), while the LPW powder is mostly martensite (α’). Particle size and morphology of the two powders are similar - spherical (roughly) particles with diameters in the range 20-60 μm.Table 2. Powder propertiesEOSLPWComp- osition (wt%)ON0.050.150.070.03+62 μm7.0%7.4%Particle size+44 μm42.7% 32.1%23 μm1.1%8.9%2-ΘAs-built materialsFigure 1: As-received powders have different phase composition. EOS powder is mostly austenite (γ) while LPW powder is mostly martensite (α').Table 3 shows the oxygen andnitrogen content of the as-built specimens. As expected, the EOS material has substantially more N2 than the LPW material. However, the gas atmosphere during processing does not have a significant effect. Nitrogen in the metal is not gained or lost significantly during SLM processing.Table 3 shows Feritscope results for as-built specimens.

Nitrogen in the metal is not gained or lost significantly during SLM processing.Table 3 shows Feritscope results for as-built specimens. The martensite content of the EOS material processed in N2 is less than 4% for both processing atmospheres, i.e. it is almost completely austenite. Processing the EOS powder in argon has only a very small effect; the material is still more than 96% austenitic. The as-built LPW material is very different. It is at least 76% martensite in either process atmosphere, comparable to the commercial, condition A materials. As nitrogen is known to be an “austenite stabilizer” (Ulyanin, 1969), it is clear that the difference in as-built phase composition is due to the difference in nitrogen content of the powders.441Table 3.

Element and phase composition for SLM processed materials.PowderEOSLPWBuild atmosphere Layer (μm)20N240Ar40N240Ar40O (wt%)0.040.040.080.07N (wt%)0.170.140.140.030.03Feritscope martensite (%)As-built 650 oC heat treat 788 oC heat treat0.70.7340.73.78276Figure 2 shows room temperature XRD patterns for specimens built (40 μm layer) from both powders using nitrogen and argon atmospheres. The R and L patterns are for the analysis plane perpendicular and parallel to the build plane, respectively. These as-built patterns are consistent with the Feritscope measurements. The LPW material is fully martensitic in all cases. The EOS material has some amount of austenite in both atmospheres, but more when built under N2.

The LPW material is fully martensitic in all cases. The EOS material has some amount of austenite in both atmospheres, but more when built under N2. The austenitic EOS material shows significant texture; the (111) reflection is very strong when the analysis plane is perpendicular to the build platform and very weak when the analysis plane is parallel to the build plane.Optical and electron micrographs (Figure 3) of N2 as-built EOS material show a uniquemicrostructure. The optical micrograph shows the macrostructure typical of SLM processed material with overlapping, bowl-shaped features that result from solidification of the melt pool created by each laser scan. At higher magnification each of these features is seen to consist of very fine, parallel, cylindrical austenite grains. The average diameter of these is approximately 0.5 μm, while the length can be 40 μm or more, sometimes continuing across the boundary between two solidification features.

The average diameter of these is approximately 0.5 μm, while the length can be 40 μm or more, sometimes continuing across the boundary between two solidification features. This strong grain orientation is consistent with the observed XRD texture and indicates that solidification occurs with crystal growth perpendicular to the close-packed (111) austenite planes.Effect of Heat TreatmentThe Feritscope measurements (Table 3) for EOS material processed in N2 show littleconversion of the austenite for the 650oC heat treatment and very substantial conversion for the heat treatment at 788oC. High temperature XRD results provide additional understanding of the effect of heat treatment temperature. Figure 4 shows the intensity of austenite (111) and martensite (110) reflections for specimens mounted so that the analysis plane is parallel to the build plane. Because of strong preferred orientation, the austenite (111) reflection is weak.

Because of strong preferred orientation, the austenite (111) reflection is weak. Heat treatment up to 700 oC has no effect on the phase composition either at high temperature or during cooling. When the heat treatment temperature is 750 or 800 oC, this reflection becomes substantially stronger and this “new” austenite transforms to martensite upon cooling. This change in behavior coincides with the α-γ transus temperature for iron (727 oC), the temperature at which ferrite transforms to austenite upon heating.442Figure 2: XRD patterns of as-built material are for the analysis plane perpendicular (R) and parallel (L) to the build plane. Consistent with Feritscope measurements LPW is mostly martensite (α') while EOS is shows significant austenite (γ).Figure 3. Optical (L) and scanning electron (R) micrographs show macrostructure typical for SLM process and strongly oriented, fine austenite grains.443Figure 4.

Optical (L) and scanning electron (R) micrographs show macrostructure typical for SLM process and strongly oriented, fine austenite grains.443Figure 4. High temperature XRD of N2-processed EOS material shows significant change in austenite characteristic at temperatures above 700 oC. This austenite transforms to martensite upon cooling to room temperature.Mechanical propertiesThe metastable austenitic stainless steel created by SLM of high nitrogen 17-4 powderhas interesting mechanical properties. Figure 6 shows the stress-strain curve for a tensile test of as-built EOS material using 20 μm layer thickness. The test specimen was built in the z-direction, i.e. the long axis is perpendicular to the build plane and parallel to the (111) planes of the oriented austenite grains. The elongation-to-fracture is almost 40%.

the long axis is perpendicular to the build plane and parallel to the (111) planes of the oriented austenite grains. The elongation-to-fracture is almost 40%. Continuous Feritscope measurement during the tensile test shows that austenite starts to transform to martensite somewhat after reaching the yield point and continues until fracture. The final, plastically deformed material is 50% martensite.The transformation from austenite to martensite also relieves residual stress in as-builtmaterial. Three rectangular parallelepipeds 5x5x50 mm were fabricated from the EOS powder under N2. Residual thermal induced stress in the beams was measure by measuring distortion of the beams after release from the build plate. Table 4 shows the results for no heat treatment and for stress relief heat treatments at 650 oC (EOS-recommended) and 800 oC. The 650 oC temperature shows little change compated to no heat treatment.

Table 4 shows the results for no heat treatment and for stress relief heat treatments at 650 oC (EOS-recommended) and 800 oC. The 650 oC temperature shows little change compated to no heat treatment. Raising the temperature to produce the phase transformation completely eliminates it.444Figure 5. As-built EOS 17-4 stainless steel exhibits very large elongation to failure due to strain-induced transformation of the metastable austenite to martensite.Figure 6. Stress relieving heat treatment is effective for temperature that produces the austenite-martensite phase transformation.ConclusionThe combination of high nitrogen content and very fine, columnar grain structure yieldsan unusual metastable austenitic form of 17-4 stainless steel. The austenitic structure is stable up to the ferrite-austenite transus temperature. Heated above this temperature the material decomposes to the typical martensitic form upon cooling to room temperature.

Heated above this temperature the material decomposes to the typical martensitic form upon cooling to room temperature. This occurs even though the nitrogen content of the metal has not changed. Thus, nitrogen content is not enough to explain the formation of stable austenite using EOS powder. The very fine, elongated grain structure also is needed to resist transformation to martensite during cooling.The metastable austenite form of 17-4 stainless steel has certain unique and useful characteristics. Plastic deformation induces austenite to martensite transformation, resulting in a very large elongation prior to fracture. While this transformation-induced-plasticity (TRIP) has been observed in other steel alloys, it is not observed in 17-4 under normal processing methods.

While this transformation-induced-plasticity (TRIP) has been observed in other steel alloys, it is not observed in 17-4 under normal processing methods. Also, transformation of the austenite by heat treatment above the transus temperature fully relieves residual stresses produced during the SLM process.More generally, this study shows that microstructure and phase compositions producedby SLM of metal alloys can be very different than those produced by conventional manufacturing methods. Perhaps, with better understanding, these differences can be exploited to produce performance benefits unique to additive manufacturing methods.Acknowledgement The authors thank Mr. John Krawczyk who performed the high temperature XRD experiments. This research was supported by Dr. Ignacio Perez through Office of Naval Research grants N00014‐09‐1‐0147 and N00014‐10‐1‐0800.445ReferencesFacchini, L., N. V. Jr., I. Lonardelli, E. Magalini, P. Robotti and A. Molinari (2010).

Microstructure control has interesting implications for other research as well, such as the MASK & DEPOSITS ap proach of forming objects.1. Rapid ToolingRapid manufacturing of tooling for injection molding, stamping, composite layup or similar pro cesses where the shape of the tool is critical is a challenging problem with considerable commercial potential [1]. The creation of such tooling by arc-spraying zinc and zinc alloys has been in the commercial literature for at least 25 years [2], and thick sprayed zinc structures have been in the lit erature for 65 years [3]. These alloy systems, however, are relatively soft and prone to wear and loading failure; their usefulness is largely limited to prototype tooling applications and low pressure applications such as reaction injection molding tooling. Tools made from ferrous systems[4, 5], as shown in Figure 1, are of far greater applicability, both for superior prototyping and limited pro duction.

Tools made from ferrous systems[4, 5], as shown in Figure 1, are of far greater applicability, both for superior prototyping and limited pro duction. They present much greater wear resistance compared to zinc systems, they are stronger, and they withstand the demands of elevated temperature service. Further, the demands of the injec tion molding application are well matched to the ferrous shell structures produced by arc-spray, particularly the support of hoop stresses in the tool structure, the tolerance of compression in the the ferrous shell, and the tool face wear resistance to abrasive plastics such as glass filled nylon.© Copyright 1991 Aluminum Company of America and Carnegie Mellon University213FUSSELL, KIRCHNER, et al:, Controlled MicrostructureFig. 1. A 420 stainless st¢elsprayedinjection mold for a flying disk.Our system uses a computer based geometric modeling system (NOODLES [6]) to describe the part an<ipatterns needed to make the Part'smoldordie.

A solidfreef0rtnfabrication process, a stere olitbogral'hy apparatust inthis ca~e,Jlutonornouslycreates the pattemina matter of one or two days. The shells are fabricated by robotiCaIlYispraying metal using an arc-spray device to create the tooling face and structure. The back side of the tooling cavity is then filled with a support material to sustain the compressive service loads. Figure 2 shows a cross section of such a tool.FrameBacking MaterialFig. 2. Schematic cross section of a sprayed tool. The frame interior is angled to help support compressive loads on the tool face.t Stereolithography has been commercialized by 3D Systems, Inc.

of Valencia, California under US Patents 4575330 B1, and 4 929 402.214FUSSELL, KIRCHNER, et al:, Controlled MicrostructureThe synergy of a coherent, computer based three-dimensional modeling system, a rapid prototyp ing fabrication device, and robOtically based spraying system has made this approach for manufac turing low volume tooling economically appealing, particularly for geometrically complex shaPeS: the time needed to manufacture such a tool is of the order of one week; and the cost of the manufactured tool is substantially less than a conventionally made one-of-a-kind mold or die.The direct linkage of the part's computer model and the part's tooling results in paperless manufac ture; changes in the part's geometry are directly communicated to the next iteration of tooling, and unacceptable aspects of the design, from a manufacturing perspective, are communicated back to the part's computer model.

Using the robOt to manipulate the arc-spray gun has three striking ad vantages: any particular schedule for spraying a mold or die half is rePeatable; moreover the robOt is precise, meaning, for example, an intended standoff of 10 cm is precisely executed; and the robOt is reprogrammable for new tooling designs. This combination yields immediate benefits in the quality of sprayed ferrous alloy shells. The robOtic repeatability is required for consistency in the sprayed shell, as well as consistency in this experimental work. The robOt, being a pro grammable mechanism with considerable freedom of motion, is also capable of spraying complex surface geometries.The basic fabrication of these sprayed ferrous tools had been problematical.

In making sprayed tooling, metal is fITst sprayed onto a substrate, and then removed from that substrate; thus the inter face layer between substrate and coating is by design weak in order to permit that separation.After the arc-spraying is completed, the shell and the substrate still contain considerable internal stress, albeit in equilibrium. These thermally induced stresses can be very large [7]. When the shell is separated from the substrate,. the shell defonns as it comes to a new equilibrium. Our current process art limits this to roughly 1 mm deflection in 500 mm length of shell for a 2 mm thick shell. Secondly, the sprayed shells are backed with a mass castable ceramic or filled epoxy to principally provide tool compressive strength. If the shell and the backing material have different coefficients of thermal expansion, then the normal thermal cycling of a tool will create shear at the shell, back ing interface.

If the shell and the backing material have different coefficients of thermal expansion, then the normal thermal cycling of a tool will create shear at the shell, back ing interface. The presence of internal stress in the shell, as well as a mismatch of coefficients of thermal expansion, also raise questions of the shell's geometric stability at elevated temperatures, and over long times. Finally, the spray process is poorly suited for spraying into narrow channels and small aSPect ratio holes.As much of the tool's function is determined by its geometry, much of the tool's strength and wear characteristic's are determined by the microstrUcture of the sprayed shell. The focus of this work is the microstrUcture of the sprayed shell,. with emphasis on designing aspects of the microstrUcture to meet the demands placed on the tool's inner shell.

The focus of this work is the microstrUcture of the sprayed shell,. with emphasis on designing aspects of the microstrUcture to meet the demands placed on the tool's inner shell. This is one important component of the re search needed to make successful ferrous tools; this work will help guide the future efforts in the larger research program.2. Metal Spraying Techniques2.1 Osprey, plasma, andflame techniques..A variety of techniques are available for depositing metal by a spray process. The Osprey process [8,9] provides large deposition rates (1000 kg hr 1) by atomizing a molten stream of liquid drawn from a pc>ol of liquid metal. This is generally done in a chamber under inert conditions. It is possible to superheat the liquid pool. The deposition rates are such that a contiguous liquid surface is present on the substrate.215FUSSELL. KIRCHNER. et al:.

It is possible to superheat the liquid pool. The deposition rates are such that a contiguous liquid surface is present on the substrate.215FUSSELL. KIRCHNER. et al:. Controlled MicrostructurePlasma systems [10-14], at the other end of the spectrum, deposit material at a rate from 0.1 kg hr 1 to 5 kg hr 1. These systems function by propelling powdered material in a stream of gas heated by an electrical arc; gas temperatures are reported to be as high as 20000 K [14]. Deposition rates of a few kg hr- l are problematical, for typically the powder particles are not en tirely melted. Plasma systems permit a wide latitude in choice of materials, as well as good of con trol of the resulting microstructure; they are mostly used for coating.Flame systems, similar in concept to oxy-acetylene cutting and heating torches, melt a powder, wire, or stick of material in the flame and then use the combustion gases to blow the molten parti cles to the substrate.

The deposition rates vary between 4 kg hr 1 and 20 kg hr 1. As this process essentially uses a heating torch, the substrate is heated, while the particles are conveyed in a gaseous stream of combustion products.D.C. voltageelectrode wirestandoff distancedeposit and substrateFig. 3. Schematic of an atc--spray gun and deposited particle flight times.High velocity variants of flame and plasma systems greatly increase the kinetic energy of the parti cles in order to reduce the porosity of the sprayed material. The high speed flow of molten material also serves as an abrasive to the substrate.2.2 Arc-spray. The effort reported here uses an arc-spray system [15, 16]. It is comparable in cost to the low cost flame system while avoiding the inherent products of combustion, it is ex~ tremely easy to use, and it uses widely available materials.216FUSSELL. KIRCHNER. et aI.:. Controlled MictostruetureThe arc....spraygunis arranged as in Figure 3.

KIRCHNER. et aI.:. Controlled MictostruetureThe arc....spraygunis arranged as in Figure 3. Two consumable electrode wires are fed through contacttips to the area of the arc. A >D.C. power supply establishes an arc between the wires, melting them in them:c.A column of atomizing gas, ranging from 480 kPato 690 kPa (70 psito 100 psi)iablates the molten materialfromthe wires,at0t11i~es thernolten droplets and carries them, in a spray,to/the substrate. For steel systenls,thearc voltage typically ranges from 26 volts to 31 volts; the arc cUlTentranges from 5~ampst030(JampsgivingatemI>-erature of 104 K in the arc [14]. Deposition rates for anarc....spraysystemrangefrom 1 kg hr- 1 to 20 kghr- 1. The struc tures resulting from the •arc-spray process are suitable •for the tooling applications at hand[16, 17], and the arc.....spray process permits a deposition rate that atlowsa timely buildup of the tool shell thickness.2.2.1 Parameters.

Within the arc-spray process, there are a number of parameters that can be controlled to affectthe quality of the deposited shelL[18, 19].

These principally include:standoff distance-the distance the particles must travel from arc to substrate;gun orientation -the orientation of the•• spray gun with respect to the substrate's surface normal, and thus the.direction of impactof the sprayed particles upon i the substrate;traverse speed -the traverse speed of the spray gun over the substrate; this di rectly affectsithe flux of particles arriving at the substrate;gas pressure -the atomization and accelerating gas pressure in the spray gun;arc energy-the energy being consumed in the arc; the resistance of the arc is essentially constant at a constantmaterial consutl1ptionrate, so the arc power (and hence the arc specific energy) i is set by the power supply voltage;material consumption rate -.arc per unit time; the amount ofrnateriatbeingpresentedto the i it is specified as a feed rate of the consumable electrode.

iNot all ofthistl1aterial is deposited -- a fair proportion is lost to overspray.Table 1 shows an overview of the relationship between these parameters and some effects in the spray and shell'snrlcrostructure.2.2.2 Solidification and Microstructure.Our aim is to influence the shell's microstructure indirectly by directly influencing the spray's characteristics. The porosity andoxide levels should be kept at a minimum (ideally zero), the particle sizes should be uniform, each with a uniform temperature, while imparting no heat load onto the substrate or previously sprayed shell. A measure of compromise is needed in setting the spray parameters.The porosity in Table 1 can be most directly influenced by either increasing the particle's kinetic energy or increasing the time for coalescence on the substrate; gas pressure, standoff distance (and therefore their speed) and adding more superheat to the particles are the means to this end.

Orientation has a ftrst order effect on porosity for shadowing reasons - previously deposited parti cles will act as obstructions to incoming particles; worse, the hole formed in the shadow of a pre-217FUSSELL, KIRCHNER, et al:, Controlled Microstruetureviously·deposited particle is large, andiscluster¢dwithothersuch.holes. Figure·4 schematically shows this effect, and Figure 16 shOwSlhe effectin the microstructure. We have found thatPorosityOxideSize of ParticlesTemp. of ParticlesHeat Load on ShellOrienta tion of LaminaeStandoff DistanceoTorch OrientationTravel SpeedGas PressureoooEnergy of ArcoooMaterial Deposi tion RateoooTable 1. Qualitative effect ofparatneters on spray characteristics. '.' impliesfiTSt order contribution, '0' implies a lesser influence.orientation angles greaterthan300 leadtQunacceptaple p<Jr0sity.

Qualitative effect ofparatneters on spray characteristics. '.' impliesfiTSt order contribution, '0' implies a lesser influence.orientation angles greaterthan300 leadtQunacceptaple p<Jr0sity. Oxide formati?nisa kineticreac tion, so,beyond eliminatingorreducin$~xygenjnthevicinity of the moltenmet8.lQr1oweringthe temperature of thereactingptoducts,t¢ducingthetimefor the formation of oxides is significant. Particle size is principally determined by atomization in the arc; coalescence in the arc and on the0o ,&iiis.. .. ...Fig. 4. A schematiC ofthe shadow effect ofporosity formation.substrate are minor secondary factors in arc-spray.

.. ...Fig. 4. A schematiC ofthe shadow effect ofporosity formation.substrate are minor secondary factors in arc-spray. Beyond the first layersd~posited, the orienta tion of the built-up laminar structure is controlled exclusively by the orientation of the torch.218FUSSELL, KIRCHNER, et al:, Controlled MicrostructureIndividual lamella morphology is a function of both the particle size, as well as the orientation of the torch with respect to the substrate's surface. The impact of the molten particles on the substrate is greatly influenced by the angle between the particle's trajectory and the local surface normal; ori entations increasingly distant from the surface normal produce a more fragmented lamella. This ef fect, combined with the shadow effect, produces undesirable microstructural features.

This ef fect, combined with the shadow effect, produces undesirable microstructural features. Narve meth ods of depositing metal shells result in poor quality shells; spraying with the intention of creating sound microstructures has led us to investigate methods of miling shells with lower oxide levels, lower porosity, oriented lamella, and shells built of sprayed composites.3. EXPE.RIMENTAL ARRANGE.MENTThe experimental arrangement used in this work is organized around a robotically moved arc-spray gun. Both the robot and the arc-spray system are coordinated by an IBM-AT computer; all of the experiments reported here were controlled from the IBM-AT.

Both the robot and the arc-spray system are coordinated by an IBM-AT computer; all of the experiments reported here were controlled from the IBM-AT. The arc-spray gun and power supply are commercial products (the gun is a Miller BP-400 and the power supply is a Miller Mogularc 400&); the arc-spray controlleris a Miller custom built process controller that is interfaced to and controlled by the external computer which sets the wire feed rate and arc voltage, as well as starts and stops the arc process. The robot is a GMF 8-700 6 axis robot, with additional degrees of free dom in a rotating table. Programs are uploaded and downloaded through the IBM-AT computer, and the IBM-AT computer initiates robotic motion. The robotic manipulator provides a high degree of motion repeatability in making.

Programs are uploaded and downloaded through the IBM-AT computer, and the IBM-AT computer initiates robotic motion. The robotic manipulator provides a high degree of motion repeatability in making. the sprayed structures; the computer controlled power supply and arc-spray system gives a similar repeatability to the process variables in the sprayed structures.Voltage vs Arc Specific Energy (420 S.S.; 1.6 mm wire)......... b.() ~ ~ 3.2e 6 i) 3.0 e 6 ~ 2.8 e 6 ~ 2.6e 6 .~ 2.4 e 6 VJ 2.2 e 6 - ~/3.8 kghr-1 ~-1.9 kghr-1 ~L~' 1 ....~ 4.7kghr-~..~~............"""~'~~­.....~ ~.....,~","''''''' .. ~<.........,..""""" ~~ ., , , 29 31 30 Voltage (volts) Fig. 5. Arc voltage vs arc specific energy with the deposition rate as parameter.

~<.........,..""""" ~~ ., , , 29 31 30 Voltage (volts) Fig. 5. Arc voltage vs arc specific energy with the deposition rate as parameter. Essentially the energy per unit mass melted is proportional to the control voltage, but independent of the deposition rate., 282732i33For a given deposition rate, the arc current and arc resistance are relatively constant as the arc volt age is varied; thus the power consumed in the arc is largely controlled by the voltage setting. Figure 5 shows the effect of increasing arc voltage on specific energy, all other variables held con stant. With this type of equipment, this is the simplest way to increase the energy available to heat the metal particles.219FUSSELL, KIRCHNER, et al:, Controlled MicrostructureThe experiments were conducted at the lower end of the arc-spray parameter ranges t with a depo sition rate of about 2.0 kg hr 1 and specific energy expenditure of 2.6 MJoules kg- 1 (0.74 kWH kg-I) of metal consumed.

With the programmed torch travel speed of 500 mm s-1 a deposit of 0.1 mm is applied during each cycle of 24 passes. In all cases, the standoff distance was the same 16.5 cm (6.5 inches) from the arc to the substrate. Similarly, the orientation was, except as noted belowt normal to the substrate (i. e. t the particles traveled along a course parallel to the sur face normal).The atomizing gas, in our facility, comes either from an air compressor (capable of supplying 1.2 MPa (175 psi) pressure) or from a cluster of compressed gas bottles. The process produces a broad spectrum of particle sizes, ranging from +100 micron diameter particles to sub-micron parti cles.Due to the large surface/volume ratio of the sprayed droplets t oxidation in flight and after im pingement is a problem.

The obvious way to circumvent the detrimental effect of oxidation and avoid the ensuing brittleness of the deposit is to minimize the partial pressure of oxygen in the gaseous atmosphere [20] and keep turbulence as low as possible. The influence of these factors was studied by choosing three experimental setups, as shown in Figure 6. Either the robotically manipulated jet was directed towards the substrate without any protective cover (Figure 6a), or it was protected by a shroud (Figure 6b) that minimized turbulence of the surrounding atmosphere. In this arrangement the hot jet and the yet hot deposit is mainly flushed by the gas of the jet. Additionally an arrangement was tried (Figure 6c) where the spray gas was flushed through the shroud to provide inert gas for the turbulent mixing of the jet. The shroud is attached to the nozzleGas~GasJGas"'.---J----.tII 1 ((a)(b)(c)Fig. 6. Shroud arrangement for inert atomization and cover gas. (a) no protective cover.

The shroud is attached to the nozzleGas~GasJGas"'.---J----.tII 1 ((a)(b)(c)Fig. 6. Shroud arrangement for inert atomization and cover gas. (a) no protective cover. (b) protective shroud. (c) protective shroud with protective gas flow.and moves with it. Although the distance between nozzle and substrate is kept as constant as pos sible in the robotic operationt there always remains a gap between shroud and substrate. Thus it is practically not possible to prevent contact between the surrounding atmosphere and the jet or the hot deposit, but obviously the shroud protects the jet rather well when and after it leaves the nozzle220FUSSELL, KIRCHNER, et al:, Controlled Microstructure(and is still slow, hot and reactive). Near the substrate turbulence might draw in some of the sur roundingatrnos~here,buttherethe temperature ofthe droplets and.. thus the rate of oxidation is al.. ready lower.

ofback~and ..forthrnotion •.across a substrate (7 nun1owcaroon steel plateinJllOstcases)Step 2. waitforthetet1.lperature ofthe sprayedsl1ellto.returnfrom its heated state (typicallyStep 3.6O°C)to some nominal level(typically 40°C) repeatuntiladequate shellthickness for sectioning.was reached..Generally this prace durewas repeatedeighttotetrtimes,giving atypical shell thickness of 1.3 mm.Various materials were sprayed onto mild steel that had been glass bead peened (compositions are in weight%):(a) Plain carbon steel 1080. The composition is 0.8 % C, remainder Fe. (b) Stainless steel 420 (a martensitic stainless steel) with a.composition 0.15% Cminimum, 1% Mn,l% Si, 12-14%Cr, 0.04%P, and 0.03%S(c) Bronze of 2%Si,.remainder Cu. (d) Invar of 36% Ni, remainder Fe.The samples sprayed in a completelyinert atmosphere were made in a closed vacuum chamber. The chamber was. evacuated to 0.7 kPa (5 torr), then backfilled to 57 kPa(430 torr) with argon.

The chamber was. evacuated to 0.7 kPa (5 torr), then backfilled to 57 kPa(430 torr) with argon. The arc--spray ••gunwasinsertedinto the.charnberanddutlngsprayingavane. vacuum pump. was used toholdthepressureat57kPa.Theatomizationgas for these experiments was argon. ·The substrate for these experiments was a 46 cmdiametermandrel.221FUSSELL.inthe "'......:..n...... periphery of the spray's cone4.2 substrate. the i'1""~""""''f'' ated arc--spray system cool ticles are cooling at least at Moreau [24] has shown cooling zirconia at 108 K s-l, _"""'n.'....... "... complete in 20 flseconds. This .:JM.lii.VlJO"lJOet al:. Corltro:Ued Mic:rostructure....""""",,:u:o.... 40 and 70 m s-l.

This figure shows that the total deposit is built up of lamella which range from 4 to 10 Jlm thickness and range from 50 to 100 Jlffi length. Each of these lamella is a solidified droplet, indicating that the average droplet-diameter ranges from 10 to 20 Jlffi in diameter to 100 Jlm. The largest lamella in these micrographs suggests an original particle of perhaps 140 Jlm in diameter.The micrograph indicates several things: droplets solidify in isolation - there has never been a liq uid pool on the surface. The uniformity of the size distribution implies that splashing is not impor tant: upon impinging on the surface, droplets hardly break up but flatten out to lamella before they solidify. The arc-gun was oriented so that the particles arrived normal to the surface for the material sprayed in this sample. Three different shades of gray are visible in Figure 8, particularly Figure 8(b), white, gray and black.

Three different shades of gray are visible in Figure 8, particularly Figure 8(b), white, gray and black. By X-ray energy dispersive analysis, Figure 9, we verified the white area as being metal, the grey areas as oxide and the black areas as pores. The original composition of Invar is Fe-36 w% Ni; the composition has been changed in this sample by the heavy oxidation (roughly 30% of this sample is oxide). The bright regions are about Fe-40 w% Ni, and the dark regions are roughly Fe-18% Ni. Nickel oxidizes more slowly than iron, so it is unsurprising that the oxide regions are predominantly iron.(a) Spectrum for bright (metal) regions(b) Spectrum for dark (oxide) regionsFig. 9. X-Ray dispersive analysis of atmosphere sprayed Invar.The morphology of the oxide particles raises the interesting question of how the formation of lamella (and not spheres) of oxide was possible.

X-Ray dispersive analysis of atmosphere sprayed Invar.The morphology of the oxide particles raises the interesting question of how the formation of lamella (and not spheres) of oxide was possible. If liquid metal droplets had hit the surface, and had splat out into lamella before becoming oxidized, one would expect the lamella to be sur rounded by oxide, giving a microstructure with intergranular oxide and, presumably, undesirable mechanical properties. The presence of oxide lamella indicates another mechanism: metal droplets have oxidized in the arc and in flight between the nozzle and the substrate, forming droplets of liq uid oxide that hit the surface and squash out to lamella. The discussion of section 4.1 indicates that the droplets spend about 3 msec in flight, they have been heated to less than the arc temperature 104 K [14], but more than the oxide liquidus (1670 K for Fe304, 1830 K for FeO).

Apparently these liquid iron droplets can oxidize to liquid Fe304, Fe203, and FeO while being kept well su perheated for 3 msec.5.2 Oxide. The oxide and porosity levels in the deposited shells are shown in Figure 10 with micrographs of these shells. The oxide and porosity measurements shown here have been per fonned using an image analysis program analyzing scanned images from photomicrographs. The224FUSSELL, KIRCHNER, et al:, Controlled Microstructurelevel of uncertainty in these measurements is high. Fowler [27].has shown that with eight repeti tions on a single sample, it is possible to operate within a 95% confidence band for porosity mea surements, but oxide measurements require more measurements. This can be seen in this data in the 0.8%C - Fe column; the oxide and porosity values vary in unexpected ways compared to the other materials. Further, spray-gun designs greatly influence the porosity and oxide levels in the sprayed shell.

Further, spray-gun designs greatly influence the porosity and oxide levels in the sprayed shell. The measurements here, therefore, should only be used to represent what trends are possible with inert gases.With the open unprotected arrangement of Figure 6a one obtains an oxide content of approximately 30% for both steels and the Invar alloy if air is used for atomization. This drops to 10% to 15% oxide for a nitrogen atomization gas in an air environment. If oxidation had occurred only in the nozzle, no oxidation should have occurred at all, but the turbulence of the air-atmosphere still causes oxidation. This is largely prevented by using the arrangements of either Figure 6b, or better still, Figure 6c. These decrease the oxide content to about 8%. If argon is used in the nozzle with the protection of Figure 6b or 6c, the oxide content falls to less than 8%.

These decrease the oxide content to about 8%. If argon is used in the nozzle with the protection of Figure 6b or 6c, the oxide content falls to less than 8%. If the experiment is con ducted in a vacuum chamber backfilled with argon, the oxide content drops to less than 2%, but this arrangement is experimentally expensive. In the argon backfill~d chamber, the measurement of oxide reflects wire contamination more than in-spray formed oxide.5.3 SEM Metallography. Figure 11 shows a backscatter image of the sprayed eutectoid steel (0.8% C). Depending upon the quench rate of the material, and the temperature history of the material, several phases and mixtures of phases are possible: pearlite, a mixture of ferrite (a) and cementite (Fe3C), upper bainite, lower bainite, martensite, and retained austenite (y) [28]. For this composition of steel, pearlite and the bainite transformations can be avoided if the cooling rate from 996 K to 473 K is at least 200 K s-1 [29].

For this composition of steel, pearlite and the bainite transformations can be avoided if the cooling rate from 996 K to 473 K is at least 200 K s-1 [29]. If the cooling rate is much faster, martensite particle formation will be suppressed as well [30, 31]. The 25 J.1m particle in the center of Figure 11 shows an already solidified particle when it arrived at the substrate. This particle is en~irely composed of retained austenite. The lamella below and surrounding the particle arrived at the substrate in a molten state, and subsequently solidified. The cooling of this lamella was such that martensite par ticles formed in the last deposited, or upper, portion of the lamella, while the lower region of the lamella is only retained austenite. This shows the cooling rate was highest near the lamella-sub strate interface, and that the region of the splat away from the interface stayed longest at high tem perature.

This shows the cooling rate was highest near the lamella-sub strate interface, and that the region of the splat away from the interface stayed longest at high tem perature. The lamellre and particles in this figure probably arrived at the substrate nearly simultane ously; the solidified particle forced itself onto the previously arrived lamella, forcing the liquid sur face up and around. In the larger lamella of Figure 11, small pores, of the order of 0.1 J.1m to 0.5 J.1m in diameter, are also visible. These appear in all of the sprayed materials from this process, and are an artifact of the violence of the ablation and atomizing processes.The microstructure in the lamella of arc-sprayed material is the same as the structure of splat-rib bon materials, while the interfaces between the lamellre are much worse. This is the problem of the process. The production of splat-ribbons is a continuous casting process.

This is the problem of the process. The production of splat-ribbons is a continuous casting process. The arc-spray process deposits liquid droplets individually on the solid surface, each of these droplets undergoes thermal stresses during cooling and this can lead to decohesion or microcracking along the lamella-lamella interface. The usual consolidation methods, like HIP or forging (cold or hot working) are in conflict with our intentions for this sprayed material, and therefore more subtle methods like laser glazing or shot peening of each layer could and should be used. For the unconsolidated deposits225FUSSELJ."K.IRCHNl3R, et al:, Controlled Microstructure420 Stainless Steel0.8% C Steel50~II1lL.I__....200x: 23-27% oxide.

10 (b) shows metallographic sections of the Copper and Invar deposits227FUSSELL,KIRC~R, et al:, Controlled MicrOStr'UctureFig. 11. SEMnUcrophotograplrofthespraYed0.8%C i.steel;.Showing martensite fonnation sUIToun<ie<i•.byretainediaustenite.Tbeimartensiteparticles are acicular in appearance in this section (probablybeingJenticular in three dimensions) and vary in size; they are of the order of 0.5 Jlm by 4 Jlm.Shown or discussed in this paper, themechanicalpropenies are controlled not by the behavior of th.eindividual lamellre, but•by the· strellgthofthe interfaces between them. The· presence of inter laxnella nUgrocracksmight be detrimental intension, but is of no concern for metal shells subjected to compression.6. COMPOSITE STRUCTURESGiven the layered and disjoint.larnel1~·.fonnedinthe arc-spraYiprocess, even when sprayed in oxygen free .environments, we•• are motivated. to <ievelop processestoicontrol themicrostrDcture without cold or hot working the entire shell.

to <ievelop processestoicontrol themicrostrDcture without cold or hot working the entire shell. As discussed in section 2.2.1. and Table 1, there are a number ofprocessparatneters that can be manipulated to affectthegarticles Jormingthe lamella. Beyondithearc-sprayprocessparameters,wefollpw two cOrnPatible directions: control the mate.. rial properties by altering the materialsideposited,andcontrolthepositionand orientation of the lamella to improve the shell's characteristics. Our goal is to tune the material in the shell so it more successfully meets·the demands ofthe application.6.1 Sprayed Composite Materials One rnicrostrDctureexperiment has beento create a stratified shell, with each sttatumbeing made. ofadistinctmaterial.Figure 12. shows a structure made of alternating/layers of 0.8% Csteelan<ilnvar. The sample has been .etche<iwith picral for 15 sec onds todarken the steel. Eachlayer is abOut 0.15mm·(0.006 inch) thick.

shows a structure made of alternating/layers of 0.8% Csteelan<ilnvar. The sample has been .etche<iwith picral for 15 sec onds todarken the steel. Eachlayer is abOut 0.15mm·(0.006 inch) thick. This •sample was .sprayed using argonatornizationgas,a shroud, and a backfill shroud gas ofargon. Figure 13 shows a de tail ofthernixing region of the two materials. There are no difficulties in the inter-layer.bonding of this material, even though perhaps 15 minutes passed between the deposition of the layers.228Fig.FUSSELLtKIRCHNERtetal: t Controlled Microstructure229FUSSELL, KIRCHNER, et al:, Controlled Microstructure"~u".",,,,'1.1 in Figure 14 was sprayed using argon atomization gas, aFig. 14. Mixed pseudo-alloy of 0.8 % C steel and Invar.The nature of this mix is entirely different from that of the layered composite. The particles arriving molten still form flat lamellre in the deposit, but the distinct layering of lamellre is replaced by a random mixture of lamelIre.

The particles arriving molten still form flat lamellre in the deposit, but the distinct layering of lamellre is replaced by a random mixture of lamelIre. In Figure 14, the particles are arriving nearly simultaneously, and the bonding of the lamella is similar to that of the layered composite, as seen in Figure 15.10 f.1mFig. 15. Detail of inter-lamella region - pseudo-alloy of 0.8 % C steel and Irtvar.230FUSSELL, KIRCHNER, et al:, Controlled MicrostructureFrom the optical micrograph, it is evident that the lamellre are either Invar or iron. Apparently droplets detach themselves individually either from the steel or the Invar wire feeding the arc, and there is little opportunity for the individual particles to mix either in flight or on the substrate; the solidification rates are sufficient on the substrate to prevent mixing there.

X-ray mapping of this deposit has verified that little mixing of the Invar and iron has occurred.6.3 Oriented Lamellae A shell composed entirely of lamellre oriented in one direction may show good resistance to stresses in that direction, but will also be weak to stresses applied in other directions. An appropriate control of the orientation of the lamella in a layer will improve the shell's characteristics; this is especially interesting when the orientations and positions of the lamella are designed for specific structural features.6.3.1 Stratified Orientations The experiment shown in Figure 16 demonstrates a series of stratum, each with a new lamella orientation. The robot was used to systematically apply the material; the lamella are rotated about ±25 degrees from the horizontal. Figure 17 shows a detail of the mixing between the strata.

The robot was used to systematically apply the material; the lamella are rotated about ±25 degrees from the horizontal. Figure 17 shows a detail of the mixing between the strata. There is mixing of lamella in this region, so there should be good vertical bonding of the strata in the composite structure.An effort was made in spraying the shell shown in Figure 18 to remove some of the porosity in the shell by mechanically shot peening each layer after it was sprayed. Shot peening also has the effect of changing the state of stress in the shell, but this is not visible in optical metallography. In sam ples that are made by metal atomized by air, the shot peening also has the effect of removing a measure of the oxide on the surface.The samples in these figures were sprayed with atmosphere atomization and no shroud.

This sttucture will not withstand comer loading with any. measure of success. As a matter of fact, the material in Figure 19 cracked during/preparation. Figure 20 shows the same. geometry, but a dif ferent spray strategy; the lamellre were oriented to flow over the corner and thus present some232FUSSELL, KIRCHNER, et al:, Controlled Microstructurestructure to withstand the loading in the corner. Judicious control of the microstructure around as perities, wedges, corners, and similar features Is essential forthe assuring good mechanical behav ior of the shell. The necessary intricatemoyementsofthe nozzle over the geometric features ofthe shell are accomplished mosteffectively by a robotic manipulator.100 IlffiFig. 19. Lamellre sprayed around cOrner.These structures were sprayed using argon atomization without a shroud, In part to highlight theiQrienta.tion of the lamellre.Fig. 20.F1owoforiented lamella'.

19. Lamellre sprayed around cOrner.These structures were sprayed using argon atomization without a shroud, In part to highlight theiQrienta.tion of the lamellre.Fig. 20.F1owoforiented lamella'. around corner.233FUSSELL, KIRCHNER, et al:, Controlled Microstructure7. CONCLUSIONSThe essence of the result of the arc-spray process is a quilt of quickly solidified lamellre laced to gether into a relatively weak structure, a three dimensional jig-saw puzzle. Traditionally, this type of quilt is soaked, and then hot and cold worked to dissolve the impurities; all memory of the porosity and grain boundary imperfections are erased in such processes.We are applying the arc-spray process to quickly and repeatably create precise shapes - our par ticular goal is to make tooling shapes for molds. Soaking and working are in conflict with our goals of speed and precision; fortunately the demands of tooling are centered around hoop stresses, compressive loading, and wear.

Soaking and working are in conflict with our goals of speed and precision; fortunately the demands of tooling are centered around hoop stresses, compressive loading, and wear. The sprayed metals produce shells that lend themselves to these demands.To create successful materials in a sprayed shell, we control the levels of oxide. That is, we permit some oxide to remain - these oxides are hard, resist wear well, and are fully supported by the metal matrix surrounding them. We further tailor the microstructure to meet geometry requirements in the shell. Finally, these shells can be composed of composite structures, both in the sense of a variety of materials either mixed together or placed in stratified layers, and in the sense of differing lamella orientations. Consistent creation of this oriented structure is most effectively created by robotically manipulating the spray device.This work is part of a more comprehensive research effort.

Consistent creation of this oriented structure is most effectively created by robotically manipulating the spray device.This work is part of a more comprehensive research effort. The next steps, as guided by this mi crostructural understanding, must focus on the mechanical behavior of these sprayed materials, paying particular attention to the loading modes and wear conditions that are present in tooling applications. Further work will also concentrate on the failure behavior of these materials in the shells under service conditions; from this information, further design of the material will improve the robustness of the tooling.

ntrolling the Quality of Laser Sintered Parts Along the Process ChainS. Rüsenberg1, R. Weiffen1, F. Knoop1, H.-J. Schmid1E-Mail: stefan.ruesenberg@upb.de; hans-joachim.schmid@upb.de M. Gessler2, H. Pfisterer2 Direct Manufacturing Research Center, University of Paderborn, Paderborn, Germany1 EOS Company, Munich, Germany21AbstractThe quality of laser sintered parts, in this work, manufactured by polymer laser sintering by using an EOSINT P395 Laser Sintering system, depends on several steps along the process chain. The first step is the characterization of the powder quality, whereas the rheological and physical investigations of nylon 12 powder are shown. By changing some important influencing factors, for example the powder ratio, the powder ageing and the moisture content, the influence on mechanical and physical properties, density and porosity, are investigated. The composition of the used powder is known.

The composition of the used powder is known. The previous process (storage conditions, etc.) as well as the laser sintering process (regarding energy density, temperature, etc.) is kept constant for the duration of this work. Regarding the post process in this work the cooling down phase is investigated as well. With an automatically blasting system it is possible to keep the post process parameters blasting distance and blasting time, constant. All of the tests will be performed using dry and conditioned test specimens.This work is showing the dependence on mechanical, rheological and physical parameters by varying important influencing factors along the laser sintering process quality chain.2IntroductionAdditive manufacturing technologies will be important for different direct manufacturing applications within several industry fields.

The study “Thinking ahead the Future of Additive Manufacturing – Analysis of Promising Industries” is giving an overview about possibilities and chances using AM technologies for diverse industries. However, the way to a serial manufacturing process will be a difficult one. Lots of parameters might be known now, but there is lots of work needed to understand them [DMRC12].An important factor will be the powder quality of the input material. At the moment the material is mixed by using a fixed powder ratio of used and virgin powder. However, the condition of the used powder is typically unknown or at least uncontrolled. There is no information about the age or the thermal loading of the used powder. The thermal loading is influenced by different parameters like number of parts, height of the building job, unpacking behavior, building temperature as well as the temperature within the removal chamber.

It is necessary to find a possibility to characterize a parameter for a powder qualification. Therefore it is important to correlate rheological and particle properties with material properties. Often mechanical properties are used as a quality criterion, but there have to be investigations of additional physical, thermal and electrical properties as well. This work is presenting the correlation between mechanical1024properties and different powder qualities adjusted by melt volume rate and revised by solution viscosity.3State of the ArtThe research in the field of ageing of nylon 12 and the different influencing factors of the Laser Sintering Process on the material are approached by different methods. A commonly used method is the DSC-analysis, with which the crystalline melting temperature as well as the crystallization temperature of plastics can be determined. The window between these temperatures determines the area in which the plastic can be processed.

The window between these temperatures determines the area in which the plastic can be processed. This window is particularly big for nylon 12, which makes it very good to be used in LS. At the department of plastics technology of the Friedrich-Alexander-University Erlangen-Nürnberg it is shown in research, using artificially aged nylon 12, that the crystal melting temperature as well as the crystallization temperature is increased for aged powder. During these examinations the factor time regarding the structural change of the material was made clear. The longer the material is exposed to thermal stress, the more the structure of the material changes.

The longer the material is exposed to thermal stress, the more the structure of the material changes. This can be seen in figure 1, where a DSC-analysis of three differently aged powders is shown [Drummer10] [Gornet02].Figure 1: DSC Analysis of nylon 12 [Drummer10]Altogether the crystalline melting temperature increases more than the crystallization temperature, which results in a wider processing window of aged powder and thus allowing larger fluctuations in the building process than with pure new powder. This fact has to be considered critically regarding the change of the viscosity for old powder, since it is increased and thus influences the processing of the powder and the quality of the parts negatively. Furthermore the smaller peak shows that the needed enthalpy of fusion strongly decreases, which can be traced back to a reduced degree of crystallinity.

Furthermore the smaller peak shows that the needed enthalpy of fusion strongly decreases, which can be traced back to a reduced degree of crystallinity. In addition the melting peak is getting wider, which can be explained by a thermally caused post polymerization or post networking of the chain molecules. The increase of the crystal melting temperature can be further explained by a growth of the spheroids. By post polymerization and post networking of the chain molecules1025these can entangle more and form bigger spheroids, which are more resistant to melting [Drummer10][PDY08].Another method to analyze aged powder is the determination of the MVR (Melt Volume Rate) respectively MFR (Melt Flow Rate). At the Rapid Prototyping Center of the University Louisville it is shown that these values decrease for aged powder. Since the MVR- and the MFR- value are inversely proportional to the viscosity, the viscosity increases for aged powder.

Since the MVR- and the MFR- value are inversely proportional to the viscosity, the viscosity increases for aged powder. These results show that through thermal stress in the LS-process a post polymerization or post networking occurs. This is supported by solution viscosity measurements. Results from these measurements are shown in figure 2. With increasing age the viscosity value of the PA12 increases. The viscosity increases for a growing chain length of the polymer molecules [Drummer10] [Gornet02].Figure 2: Solution viscosity measurement [Drummer10]In order to further examine the structural change in the nylon 12, SEC-analyses have been executed. With SEC-analyses a molecular weight distribution can be determined. In addition the molecular weight distribution of sintered parts can be determined, since for the SEC-analyses the powder, respectively the parts have to be dissolved.

In addition the molecular weight distribution of sintered parts can be determined, since for the SEC-analyses the powder, respectively the parts have to be dissolved. Results of such an analysis can be seen in figure 3.The blue curve represents virgin powder and the red curve old powder, passed the laser sintering process twice. It can be clearly seen that the red curve is shifted to the right and gets wider. Regarding the structure of the chain molecules this means that the chain generally lengthens and thus the molecular mass of the molecules increases.

Regarding the structure of the chain molecules this means that the chain generally lengthens and thus the molecular mass of the molecules increases. Comparing these results with a SEC-analysis of a sintered part, an even higher increase of the molecular mass can be detected, which is logical regarding the stronger thermal stress during the process [ZHKV06] [DMRC12].1026___ Virgin Powder___ Used Powder (2x passed))M g o l ( /WMolecular Weight / g/molFigure 3: molecular weight distribution from SEC-analysis [DMRC12]Even though the viscosity of the melted plastic has a strong influence on the quality of the product, only few examinations have been conducted in the field of rheology. Using melt curves the viscosity of the melt can be displayed by means of different shear rates. Some melt curves have been determined using high pressure capillary viscometers.

Using melt curves the viscosity of the melt can be displayed by means of different shear rates. Some melt curves have been determined using high pressure capillary viscometers. The melt curves have to be seen very critical, since there is almost no process related shearing during the LS process and high pressure capillary viscometers cannot operate at small shear rates. The department of materials of the University of Loughborough has thus determined melt curves with a rotation rheometer, since this method allows the measurement of very small shear rates. However only melt curves of new and old powder have been compared to each other, in which case the differences were traced back to the post polymerization in the material [Haworth11].The rapid prototyping center of the University of Louisville has examined the change of the mechanical properties elongation at break and tensile strength with differently aged powders.

Here the part cake and the powder from the overflow container from the first build job were homogenized again and used for the next build job. This process was repeated several times. The results are displayed in figure 4 [Gornet02]. It can be seen that the elongation at break already increases for the second build job (from ~2200 PSI up to ~3600 PSI), while the tensile strength stays constant and only decreases after the fifth build job. These results, given in Figure 4, illustrate that through the mixing of virgin and old powder mechanical properties about ~3600 PSI can be achieved, which is about 63% higher than for virgin powder [Gornet02].1027Figure 4: change of the mechanical properties tensile strength and elongation [Gornet02]The comparison of all the previous results shows that the change of properties of the powder can be traced back to a post polymerization respectively networking of the molecules, characterized by various methods.

These results are unexpected with regards to the standardized injection molding process, since examinations of the structural change of plastics during injection molding show a chain reduction. But what is missing for nearly all examinations is a reference to the properties and the quality of parts that were manufactured with aged powder. The objective has to be to follow up on the present results and methods, so that they can be used as a receiving respectively quality control, in order to be able to characterize the quality of the final products.4BasicsMelt Volume RateThe basic rheological characteristic values can be described easily using the two plate model.The two plate model shall reflect the load situation that a plastic is exposed to during processing. One condition is that the plastic is adhesive at both plates and that laminar flow connections dominate. The upper plate is moved with the shear force F against the lower plate.

One condition is that the plastic is adhesive at both plates and that laminar flow connections dominate. The upper plate is moved with the shear force F against the lower plate. The lower plate is fixed, thus has a speed of v=0. The plastic in between both plates is being deformed and subjected to a shear load. The force F over the height h causes a shear stress, which is defined as follows:Since the upper plate is moved with the force F and the lower plate is fixed, a velocity profile as shown in the figure below is set. The resulting liquid layers are displaced contrary to each other.1028(1)F:shear forcev:speedh:distance between platesFigure 5:Velocity profile in the two plate modelThe effective shear stress deforms each individual liquid layer with a specific shear velocity.

The shear velocity is also known as shear rate and is defined as follows:(2)The ratio between shear stress and shear velocity is a material constant for an ideal viscous fluid at a constant temperature. This proportionality factor is called viscosity and is a material related value.(3)The viscosity is a material factor for the flow rate of a material. When flowing, the molecules of a material are displaced against to each other. The inner resistance against the friction can be described as the viscosity [PahlL95] [RH10] [Schoe11].Solution ViscosityWith a low pressure capillary viscometer the viscosity of a laminar flowing fluid can be determined according to DIN EN ISO 1624.

The basis for the measurement of the viscosity is the law according to Hagen-Poiseuille [1624].(4)Nomenclature: r = radius of the capillary, g = force of gravity, h = height of the flowing fluid, ρ = density of the fluid, = liquid volume, which flows through the capillary per time unit, l = length of the capillaryIn order to characterize polymers using the solution viscosity, they are dissolved in a solvent. The actual measurement is a comparison measurement of the polymer solution and the pure solvent. For this purpose the relative viscosity is defined by the quotient from the viscosity of the polymer solution to the viscosity of the solvent.

For this purpose the relative viscosity is defined by the quotient from the viscosity of the polymer solution to the viscosity of the solvent. All factors that refer to the pure solvent are marked with a 0 in the index [MC05].1029(5)In order to be able to better show the change of the viscosity, the specific viscosity is defined.(6)From the reduced viscosity respectively the viscosity number can be determined; whose limiting value for smaller concentrations is the intrinsic viscosity respectively the boundary viscosity [PC96].(7)For diluted solvents, which have a concentration c of less than 10g per liter, the densities of the polymer solutions and the solvent are nearly identical.

For that reason the density is canceled from all the above formulas, so that the boundary viscosity can be determined via the time measurement [MC05].(8)The determination of the boundary viscosity is done with a graphic method, with which the boundary viscosity can be extrapolated over a linearization. The following formulas are possible:(9)(10)Formula (9) shows the plotting according to Huggins and formula (10) according to Schulz- Blaschke [PC96].Exposure strategy - Part Property Profiles developed by EOS CompanyEOS Company offers standardized exposure parameters for all types of laser sintering machines. The aim of this standardization is the determination of resilient values for material properties [INT1]. By using the same parameters and running the same systems reproducible properties shall be possible. This work is performing material tests using an EOSINT P395 laser sintering machine.

By using the same parameters and running the same systems reproducible properties shall be possible. This work is performing material tests using an EOSINT P395 laser sintering machine. All kind of material properties like mechanical, thermal, electrical and physical properties are considered. Further on it is possible to test these properties by using self-created exposure strategies which is possible using the Developer Kit.

Further on it is possible to test these properties by using self-created exposure strategies which is possible using the Developer Kit. In this work the pretests using a layer thickness of 120µm and the EOS Part Property Profiles for this layer thickness are shown.As an exposure strategy the following EOS Part Property Profiles are used in this work [INT1]:ParameterLayer Thickness / µmTemperature /°CRecoater blade geometryTop Speed180181TriangleSpeed150180TriangleBalance120179Round1030Performance100178RoundTop Quality60177FlatDevelopers KitPersonalized ParameterTable 1: exposure strategy by EOS Part Property Profiles [INT1]5Experimental SetupBuilding JobA building job has been designed in order to get information about mechanical properties of laser sintered parts. Four warpage plates at the bottom are placed to get information about the warping and shrinkage.

Four warpage plates at the bottom are placed to get information about the warping and shrinkage. These test specimens have two advantages: on the one hand the warpage as well as the shrinkage of laser sintered parts can be determined. On the other hand these plates have a huge surface area and therefor it should be possible to see sink marks, which are dependent on the powder quality, e.g. the powder ageing. Further on a 5x5 matrix with tensile bars placed in z- direction is built on top of the warpage plates. The z-direction is chosen because it is the weakest direction in laser sintering [RUES11]. For testing the mechanical properties the nine central placed test specimens are used to create the average value.

For testing the mechanical properties the nine central placed test specimens are used to create the average value. At that position the assumption of a homogenous building area can be accepted.Figure 6: Test job to determine powder quality as a function of mechanical propertiesThe test specimens are placed in x-direction to avoid big edges along the recoater and to reduce the probability of any job terminations.1031Material AgeingThe tests during this work were performed by using nylon 12 cycle-powders (PA2200) from one batch.

The powder age is known, so for the 1st experiments a powder ratio of virgin powder and used powder is used with the amounts adjusted by controlling the MVR (see figure 7).l e v e lc i s a bs t n e mi r e p x el e v e lt s 1s t n e mi r e p x el e v e ld n 2s t n e mi r e p x eFigure 7: Powder usage for laser sintering processThe used powder is taken from first tests at the DMRC performed by using only virgin powder, so it is known that the used powder has run one time through the laser sintering machine: a onetime thermally charged powder (1x used powder) with a good quality is used. For the MVR 2nd experiments this powder is homogenized and mixed with a virgin powder again: a one-two time thermally charged powder (1x+2x used powder) is used in this case.Sample preparation for MVR measurementThe material investigated in this work is prepared by using a standard concrete mixer.

The mixing procedure takes 10 minutes clockwise and 10 minutes anticlockwise.As a preparation for the measurement of the MVR value it has to be ensured that no segregation of the powder has taken place in the current storage form for the nylon 12. Since this inevitably happens it is recommended to split the powder using a Retsch spinning sample divider. Eight circularly mounted glass containers are attached a rotating device. The powder is continuously fed and thus homogeneously distributed among the eight containers.For using the test sample in an MVR measurement device it is important to dry the material. The DMRC drying process is performed by using a Thermo scientific furnace. The sample is heated up to 105°C for 10 minutes. The temperature is further risen up at 140°C for another 2 minutes whereas the rising up procedure takes another 5 minutes. After the preheating process the sample1032is given to the melt volume rate measurement device.

After the preheating process the sample1032is given to the melt volume rate measurement device. To guarantee a high reproducibility and to guarantee a dry sample (no moisture adsorption on the way from the preheating process to the MVR measurement device) this procedure is treated within a time interval of 20…30 seconds. The temperatures are similar to the EOS drying process, but they are using a Sartorius MA100 moisture balance. EOS Company is realizing a rising up time from 105°C up to 140°C which is about one minute. [EOS11]Sample preparation for Solution Viscosity MeasurementPrevious to each measurement a stock solution with a concentration of 10g/L was prepared. Each series of measurements consists apart from the stock solution of four additional solutions, whose concentrations were set by diluting the stock solution. The set concentrations were 1g/L, 2 g/L, 4 g/L und 6 g/L.

The set concentrations were 1g/L, 2 g/L, 4 g/L und 6 g/L. With the weight balance AT261 Delta Ranger from the company Mettler 500±0,1 mg were placed in a 50ml Erlenmeyer flask, with an accuracy of one tenth of a milligram.Afterwards 50±0,1 ml were pipetted into the Erlenmeyer flask. During the pipetting it is important to take a reading in the lower half of the meniscus.The closed Erlenmeyer flask is placed in an ultrasound basis for 60 minutes at 30°C in order to accelerate the preparation of the solution. It is important that the temperature in the ultrasound basis doesn’t increase above the glass transition temperature of the nylon 12 (ca. 45°C), since otherwise the chain structure could be influenced before the measurement.

45°C), since otherwise the chain structure could be influenced before the measurement. Afterwards the stock solution is stored at room temperature for 24 hours, since only after this time the nylon 12 will be completely dissolved.Conditioning and cooling down phaseThe test specimen are tested in a dry condition, which means that the samples are vacuumed after the post process to reduce any influence of the environment, and at nearly standard atmosphere (TR=22°C ±1°C and rH=50%±10%). It is possible to investigate different cooling down phases and their influence on mechanical properties. All test specimens are tested dry. Directly after finishing the LS process the building chamber remains 10 hours within inert gas atmosphere within the LS machine. After this time the test specimens are unpacked immediately, respectively after duration of another 24 hours.

After this time the test specimens are unpacked immediately, respectively after duration of another 24 hours. All test specimens are vacuumed to perform tests with same conditioned samples.6Applied Testing MethodsTensile Testing - Mechanical investigationBecause there is no adequate standard for tensile bars manufactured by Laser Sinter processing, the test method standard is based on the injection molding standard DIN EN ISO 527-1 till DIN EN ISO 527-5. The test speed was 1mm/min to get information about the Young’s modulus and 50mm/min to get information about the tensile strength and the elongation at break. The velocities are given in the usual polyamide standard DIN EN ISO 1874 [527-1-5] [1874].Melt Volume Rate1033The Melt Volume Rate (MVR) is used for simple quality control and to be able define the volume flow characteristics of polymer melts during processing. The device used to determine the MVR is a low-pressure capillary viscometer.

The device used to determine the MVR is a low-pressure capillary viscometer. It stands out because the power of the piston is generated using its weight. Thus the device from this test is a pressure operated viscometer. In this work a Zwick/Roell “mflow” measurement device is used for DMRC measurement [RH10]. The melt volume rate measurement is performing by using the following parameter:Test parameter for EOS and DMRCVirgin & Used powderNominal load5 kgSample quantity4,0 gPreheating time300 sTemperature235 °CPiston distance14,6 mmTable 2: Parameter for MVR measurementThese parameters are the same for the testing with DMRC equipment as well as with EOS equipment. As a result a comparison between EOS measurement and DMRC measurement is given as well.

As a result a comparison between EOS measurement and DMRC measurement is given as well. The EOS measurement is performed by using a MeltFlow @on by Karg Company [INT1].Solution ViscosityAfter the complete dissolving of the material the stock solution is filtered, in order to ensure that no solid particles remain in the solution, which could possibly clog the capillaries. According to DIN EN ISO 307 a glass filter with the porosity 2 and a pore width of 40-100 µm is used. In order to clean the filter it is flushed under pressure with methanol several times from both sides. Methanol is a good cleaning agent since in it the m-Cresol dissolves.

In order to clean the filter it is flushed under pressure with methanol several times from both sides. Methanol is a good cleaning agent since in it the m-Cresol dissolves. All used vessels are cleaned several times with methanol after their use and left in the vent until the methanol has completely evaporated [307].1034For different measurements with different concentrations the following concentrations are created:Concentration / g/LConcentration / g/mLSolid material in 20 ml / gAmount of stock solution / mlAmount of solvent / ml100.0160.0060.1212840.0040.0881220.0020.0441610.0010.02218Table 3: Mixture for concentration of the measurement pointsWith the different concentrations it is possible to plot a curve for the solution viscosity and to detect the intrinsic viscosity as well as the viscosity number.7ResultsMelt Volume RateTable 4 shows the results for the melt volume rate measurement.

The left column is showing the MVR number. In this context the value MVR0 means homogenized used powder, which has one through the laser sintering machine (1x used powder) one time. The value MVR6 is showing the results for 100% virgin powder from one batch. The values MVR1…MVR5 are obtained for different powder qualities adjusted by using the MVR measurement. The columns MVR Value EOS and MVR Value DMRC are showing the results for the DMRC and the EOS measurement processes, respectively. It can be clearly recognized that the absolute values are different between the processes. The reasons for the differences are the different MVR measurement devices as well as the different preheating processes by using different preheating devices. However, the deviation between EOS and DMRC measurement are similar and almost constant with the MVR values for the DMRC measurement being 4.5-5.5cm3/10min higher than the respective values obtainedby EOS Company.

The MVR value for the MVR6 measurement is given as an average value of 20 individual measurements to get information about the resilience. The other values are obtained as an average from 5 individual measurements. The values are showing a huge reproducibility within a small significant range. In the right side of table 4 (2nd level experiments) the MVR values for the 2-pass powder experiments are given, using differently aged powder as explained above. It is shown that we could obtain mixtures with similar MVR values for both types of used powders. Further on the viscosity is determined by using the formulas presented above.In the next figure the reproducibility of the melt volume rate measurement is shown. A quadratic curve characterizes the powder quality for regarding different melt volume rates.

A quadratic curve characterizes the powder quality for regarding different melt volume rates. The MVR are represented for 1st level (MVR_1x-passed) and 2nd level (MVR_2x-passed) powder qualities.1036Figure 8: MVR measurement – reproducibilityThe quadratic curve is almost coinciding for both measurements. The measurements are performed with a high reproducibility and all values are reproducible with a deviation about ±2%. However, the MVR measurement is a simplifying method to get information about the viscosity. The viscosity behaves inversely which is represented for the curves Viscosity_1x-passed and Viscosity_2x-passed. The lower the melt volume rate is the higher is the powder melt viscosity. The lower the amount of virgin powder is the higher is the viscosity.Unfortunately, the MVR gives only one value at the flow curve.

The lower the amount of virgin powder is the higher is the viscosity.Unfortunately, the MVR gives only one value at the flow curve. Besides conventional methods as capillary rheometry or rotational rheometry (Haworth, Hopkinson, Hitt and Vasquez are showing differences by using virgin and used powder regarding both methods [HHHV11]), a reasonable approach to polymer properties is the detection of the solution viscosity. The next figure shows the solution viscosity with respect to polymer concentration using m-cresol as solvent (nylon 12 is soluble in m-Cresol). Different linear curves are given for chosen sample compositions. The curve is extrapolated to the y-axis where the intrinsic viscosity is given. Another regarded parameter is the viscosity number at a concentration of 10g/L. It is noticeable, that the viscosity is higher, the lower the amount of virgin powder.

Another regarded parameter is the viscosity number at a concentration of 10g/L. It is noticeable, that the viscosity is higher, the lower the amount of virgin powder. Further on the differences regarding the intrinsic viscosity aren’t as high as regarding the viscosity number for concentrations (amount of powder/amount of solution) about c=1g/cl up to c=10g/cl.1037MVR0 MVR2 MVR3_2 MVR5MVR1MVR3_1MVR4 MVR6Viscosity number at c=10g/LIntrinsic viscosity number (extrapolated)Figure 9: Solution Viscosity measurementRegarding the single values the discussions above are apparent. Using a sample of used powder (intrinsic viscosity = 149.62 mL/g), the intrinsic viscosity is much higher than with a powder sample of virgin powder (intrinsic viscosity number =93.17 mL/g) and all values are within a range of about 60 mL/g.

Regarding the viscosity number at a concentration of 10g/l all values are within an range of 110 mL/g.Furthermore it is important to observe the curve slope: The higher the slope, the higher the amount of used powder by using the same powder quality. Regarding the virgin powder it is noticeable, that this slope is the lowest one of all observed samples, except the curve for MVR4. Because of not having enough powder for the performing of two jobs, there is 2-3kg of several times used powder, taken from the DMRC cycle powder, within the powder mixture for MVR4. This powder is running through the laser sintering process much more than 2 times, so the used powder is much older than the used powder usually used in this work. It is important to know, because of the low slope for that curve. It seems to have a huge influence on the curve and on the viscosity number as well as a smaller, but visible influence on the intrinsic viscosity number.

It seems to have a huge influence on the curve and on the viscosity number as well as a smaller, but visible influence on the intrinsic viscosity number. Because of these results the solution viscosity, the viscosity number and the slope (slope = 3.23, much smaller than all other slopes) seem to be a really sensitive method to detect changes of the powder quality. Further on the reproducibility regarding the solution viscosity is tested. The powder quality MVR3 is tested two times (3_1 and 3_2 and it is shown, that there aren’t any significant deviations between the two measurements.

Further on the influence of different cooling down phases are investigated as well. The next diagram is showing the tensile strength for accomplished tensile tests using the test job explained in figure 5. The primary y-axis shows the tensile strength in MPa or the Elongation at break in %, respectively, the secondary axis gives the melt volume rate. The sample nomenclature on the abscissa denotes the adjusted MVR value (see Table 4) by the first number. The second number denotes the amount of repetitions.Figure 10: Tensile strength as a function of powder with different MVR values and a cooling down phase of 10h1039For the tensile strength after a cooling down phase of 10 hours it is difficult to find significant deviations. There are differences regarding the results for the MVR1-powder quality and the other results. The tensile strength for 2_03 seems to be the highest value which differs from the other results for 2_01 and 2_02.

The tensile strength for 2_03 seems to be the highest value which differs from the other results for 2_01 and 2_02. For the MVR3 values the results are reproducible and the total results are similar to the other results obtained from different powder qualities.The results for the tensile strength after a cooling down phase of 10 hours + 24 hours are similar. Further on there seem to be differences between the different powder qualities. The MVR3 powder seems to have the highest absolute values. This was reproduced for the other values.Figure 11: Tensile strength as a function of powder with different MVR values and a cooling down phase of 10h+24hThe absolute values are slightly higher than values after a shortened cooling down phase. This behavior is due to the laser sinter material. Semi crystalline thermoplastics are composed of crystalline and amorphous regions. The unpacking temperature using a 10 hour + 24 hour cooling down time is about 50°C.

Semi crystalline thermoplastics are composed of crystalline and amorphous regions. The unpacking temperature using a 10 hour + 24 hour cooling down time is about 50°C. The unpacking temperature for the experiments using a 10 hour cooling down phase is about 120°C and therefore much higher than the glass transition temperature, which is about 45°C. That implies that the crystallization process of the material structure is not finished. The part will be quenched and the crystallization level as well as the crystallization structure depends significantly on the cooling down process. A faster cooling down phase prefers smaller crystalline regions and, therefore, a smaller crystallization level, a smaller crystallinity. However, decreasing the crystallinity means the reduction of tensile strength, but an increase of the elongation at break.

However, decreasing the crystallinity means the reduction of tensile strength, but an increase of the elongation at break. This phenomenon is visible in the diagrams showing the tensile strength as well as the elongation at break [SCHOE11].The values for the elongation at break are high for test specimens placed in z-direction. The elongation at break for some test specimens is rising up to 20%. But it is visible that this is not reproducible. There are significant deviations although the test specimens are built up using the same exposure parameters and the same powder quality. The reason is the unpacking procedures. Depending on the person it is possible that different unpacking behaviors have an influence on the crystalline regions.

The reason is the unpacking procedures. Depending on the person it is possible that different unpacking behaviors have an influence on the crystalline regions. However, it is possible to get results about 47MPa±2 for the elongation at break in z-direction, but the reproducibility is not given.1040Figure 12: Elongation at break as a function of powder with different MVR values and a cooling down phase of 10hThe elongation at break for a cooling down phase of 10 hours + 24 hours shows a different behavior. The absolute values are smaller than after a cooling down phase of only 10 hours, but the reproducibility is shown. Further on the elongation at break is rising up together with the different MVR powder qualities and therefor with the amount of new powder. This phenomenon is visible for the first three values. For the MVR4 and MVR5 powder it seems to be a decrease.

This phenomenon is visible for the first three values. For the MVR4 and MVR5 powder it seems to be a decrease. One possible explanation might be the quality of the used powder, which has a good quality (maximum times of passing the laser sintering process: 2x). Similar tests shall be performed for typical cycle-powder to find a correlation about similar powder qualities adjusted by the MVR value, but as a mixture of different aged used powder.Figure 13: Elongation at break as a function of powder with different MVR values and a cooling down phase of 10h+24hOne noticeable result not reflected in mechanical strength is the surface quality of laser sintered parts. The surface looks really bad using a powder quality with a small amount of virgin powder like the powder mixtures MVR1 and MVR2. Most of the test specimens, placed within the1041corners of the building platform, are looking similar.

Most of the test specimens, placed within the1041corners of the building platform, are looking similar. An extract figure (figure below) shows a tensile test specimen.Figure 14: deformed tensile test specimen built with a powder quality MVR1 or MVR2, placed in a corner of the building platform.After these results there an influence of different powder qualities on the mechanical properties can be seen. It might be possible that the mechanical properties are robust toward different powder qualities using the melt volume rate as well as the solutions viscosity as a quality criterion, but there are differences regarding the surface quality as well.8Summary and OutlookThe results are showing a possible way to measure the rheological properties of nylon 12 powders. It is possible to get reproducible values using the MVR method. Further on the thermal ageing has a huge influence on the viscosity.

It is possible to get reproducible values using the MVR method. Further on the thermal ageing has a huge influence on the viscosity. The higher the amount of used powder, the lower is the melt volume rate, the higher is the viscosity. Similar statements can be made regarding the solution viscosity. It is possible to reproduce solution viscosity results within a small significant deviation. Further on the intrinsic viscosity as well as the viscosity numbers are showing distinct differences with respect to powder mixture ratio regarding the ageing of used powder. Besides the slope of the solution viscosity seems to be an important parameter as well. Depending on the age of the used powder it is possible to get a powder mixture with the same melt volume rate but with different slopes for the solution viscosity. The mechanical properties seem to be robust by investigating different powder qualities.

The mechanical properties seem to be robust by investigating different powder qualities. In some cases there are small deviations regarding tensile strength as well as elongation at break, but there are more tests needed. Further on there is a huge influence on the surface quality. Mechanical properties and surface quality have to be considered simultaneously because there will be a correlation between the above mentioned properties and the powder quality and the usage of different aged used powder. Another criterion is the quality of the used powder, which might be unusually high in the experiments presented here. Another experiment to do would be the usage of a heavily used nylon 12 material mixed with virgin one.Further on it is noticeable that the unpacking behaviors as well as the cooling down phase have a very severe impact on the mechanical properties. The results are reproducible using a longer cooling down phase.

The results are reproducible using a longer cooling down phase. It is important to know the unpacking temperature and to avoid a too fast cooling down phase.After all it is important to know the complete laser sintering quality chain, because lots of the different steps have different influence on the final quality of laser sintered parts. The next steps include a quality job with test specimen to test different characteristics like thermal, electrical and physical properties, as well as the linear and elastic mechanical properties. The influence of different powder ageing shall be tested as well.1042Figure 15: Laser Sintering Quality Process ChainAll properties shall be investigated by using all available layer thicknesses. Therefore all known influencing factors have to be considered, starting with the data preparation. It is important to know the number and position of hatch, edge and contour lines and the shrinkage has to be taken into consideration as well.

It is important to know the number and position of hatch, edge and contour lines and the shrinkage has to be taken into consideration as well. Pretests already gave valuable information about the dependence on placement, orientation and positioning, thereby it is possible to reduce the amount of test specimen within the quality job. The powder mixtures of virgin powder with differently aged recycled material, adjusted by equal melt volume rates, will be investigated as well. Further on there will be lots of other rheological investigations to get information about a sensible way to characterize the quality of starting material, because for laser sintering as a serial production method it is tremendously important to reliably characterize the input material. Next to usual nylon 12 material another, at the moment not mentioned material, will be tested. The pre-process, including storage conditions and preheating time, will be kept constant for the duration of the test phase.

The pre-process, including storage conditions and preheating time, will be kept constant for the duration of the test phase. By investigating the influence of different layer thicknesses, the EOS part property profiles shall be used as well as own developed exposure strategies. An automatic blasting system will be installed in order to reduce the influence of the blasting post process. The cooling down phase will be constant at 10 hours within the laser sintering machine and another 24 hours at standard atmosphere.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCONTROLLING THERMAL EXPANSION WITH LATTICE STRUCTURES USING LASER POWDER BED FUSIONS.S. Milward1, H. Swygart2, L. Eccles2, S.G.R. Brown1, N. P. Lavery11College of Engineering, Swansea University Bay Campus, Wales, SA20AB 2Qioptiq, Glascoed Road, Saint Asaph, LL17 0LLAbstractTuning the Co-efficient of Thermal Expansion (CTE) of a component is traditionally limited by material choice. Laser Powder Bed Fusion (LPBF) enables the designer to create complex geometries including lattice structures. When combined with a secondary material, these metallic lattice structures can be designed to exhibit different CTE’s whilst retaining stiffness. This allows the designer the freedom to adjust the CTE by changing CAD variables such as lattice angle, and member thicknesses.

This allows the designer the freedom to adjust the CTE by changing CAD variables such as lattice angle, and member thicknesses. This paper aims to develop an arrangement for CTE matched components for high precision optical systems. Development pursued using a Static Thermo-Structural Finite Element Analysis model to determine the best arrangements for the required CTE change. The results are incorporated into a new design prototype of a full cylindrical lens system in metal on a Laser Powder Bed Fusion machine.IntroductionAdditive Layer Manufacturing (ALM) has provided companies with an ability to quickly create functional parts with a range of materials from metals to plastics. Companies are increasingly looking into new ways to save money on tool up costs for low volume components. This makes ALM a very attractive proposition for the industry. Alongside the obvious benefits of ALM, comes its ability to create lattice structures.

This makes ALM a very attractive proposition for the industry. Alongside the obvious benefits of ALM, comes its ability to create lattice structures. Typical lattice structures used in the industry today consist of standardised arrangements of struts which are mostly used to fill in voids or in place of solid material. This helps save on weight and material costs, and can be generated by most software in the commercial and industrial markets. They are beneficial for an industry which takes weight reduction seriously, such as the aerospace, defence, and automotive industry. In 2017 Thales Alenia used metal ALM to produce 79 metal parts for Tlkon 3S, SGDC, and KOREASAT-8. Many other companies have also created parts for satellites, aeroplanes and rockets [1]. In 2011 Astrium created titanium EBM parts for the Atlantic Bird 7 satellite which saved 30% mass compared to traditional manufacturing means.

In 2011 Astrium created titanium EBM parts for the Atlantic Bird 7 satellite which saved 30% mass compared to traditional manufacturing means. This was achieved using lattice infill structures and topology optimisation techniques that lend themselves to ALM [2]. Additional functions can be added to play an active role in improving the characteristics of the part. These can range from the obvious such as weight reduction already actively used, to the more novel uses such as; impact absorption, thermal insulation, heat dissipation, and thermal expansion control. These have all been proven possible with unique geometries, which could lend themselves to manufacture and incorporation into existing ALM components [3][4][5].

These have all been proven possible with unique geometries, which could lend themselves to manufacture and incorporation into existing ALM components [3][4][5]. This paper will look at how parts can be designed to provide specifically thermal expansion control, by utilising different lattice designs and ALM techniques.Thermal expansion of components in an assembly is a design problem that has to be considered by most engineering industries. Issues arise when components experience a range of temperatures causing expansion and contraction, affecting its operational behaviour. Particular care has to be given when assembling multiple material parts due to different CTE values. These can cause expansion gaps, contact stresses, cycling fatigue, and component misalignment. The sponsor company Qioptiq has initiated a research program with the aim of improving the optical stability of their current and future systems.

The sponsor company Qioptiq has initiated a research program with the aim of improving the optical stability of their current and future systems. Qioptiq’s components must withstand changes in temperature from -50°C to +70°C, all whilst maintaining strength, alignment, and a lightweight design. Particular care has to be given to products which enter space. In this environment, rapid changes of temperature occur without the possibility of cooling by convection. This can lead to severe thermal gradients, stresses, and distortion. Problems can occur in satellites where optical systems 2139may be present which require dimensional stability, or antennas where a low CTE is required for a precisely tuned waveband. Space applications are a future application for Qioptiq if the concept is proven successful. LPBF has been tested with low expansion alloys such as INVAR which exhibit a CTE in the region of 1.44x10- 6(1/K).

LPBF has been tested with low expansion alloys such as INVAR which exhibit a CTE in the region of 1.44x10- 6(1/K). Due to the materials high density, topology optimisation can be used to reduce its mass, increasing its viability over traditional methods.There are several methods for controlling the CTE of a component. The more traditional methods involve material composition adjustments. This can be achieved in INVAR by varying the percentage of nickel from 36% to 56% to obtain the required CTE of between 1-11x10-6(1/K). Additional elements can also be added to INVAR to change its CTE. Some variations on INVAR include KOVAR, a FeNiCo alloy designed to exhibit similar CTE’s to that of borosilicate glass, useful in optical systems for applications such as satellites [6][7].

Some variations on INVAR include KOVAR, a FeNiCo alloy designed to exhibit similar CTE’s to that of borosilicate glass, useful in optical systems for applications such as satellites [6][7]. Variations in nickel percentage can affect mechanical properties which may be undesirable [8].A new method is proposed which utilises the unique advantages of ALM to control the thermal expansion of an annulus surrounding a glass or germanium lens. This will allow Qioptiq to control the CTE between the lens material and its metal casing. In order to achieve this goal a unique structure will be developed, comprising two materials each exhibiting different coefficients of thermal expansion. The unique concept will allow Qioptiq’s engineers to change the CTE in the design phase in CAD, by adjusting key dimensions of a bi-material lattice structure. One material must exhibit a higher CTE than the other to achieve an overall change in CTE.

One material must exhibit a higher CTE than the other to achieve an overall change in CTE. Integration into Qioptiq’s existing designs would be advantageous, however tailored new designs for unique purposes are also being considered.Figure 1) Image showing one example of an existing optical system supplied by Qioptiq.LPBF solutionsWhen using LPBF only one material can be manufactured at a time, meaning two parts have to be assembled after manufacture. This design uses a single material lattice structure in combination with a secondary outer restraining ring from a lower CTE material such as INVAR. The inner lattice material will be shrink-cooled and inserted inside the INVAR ring. This solution is capable of creating a negative CTE when performing a Static Thermo-Structural FE Simulation with Titanium and INVAR over a temperature range of 100C.2140Figure 2- a) CAD Drawing showing a shrink-fitted outer low CTE ring with a higher CTE lattice structure.

b) Von-Mises stress plot showing max stress distribution and a visual representation of deformation directions exaggerated by 300 times. C) Ti64Al manufactured with 45mm outer radius.Single print solutionsPrinting two materials together as one allows for the removal of the shrink fitting stage. It provides a greater level of design freedom where 3D structures can also be made, which opens up the possibility to control CTE in the third dimension.A Direct Energy Deposition (DED) system such as a LENS® machine could be used to print the secondary lower CTE material in situ with the lattice structure. This system allows for the creation of bi-metallic structures using two blown powders simultaneously. This can allow for replication of 2D and 3D zero and negative CTE lattice structures such as those proposed by, Sigmund et al. 1996 [9], and Steeves et al 2007 [10] [11].Figure 3a) Lattice structures proposed by Steeves showing high and low CTE constituents [10].

1996 [9], and Steeves et al 2007 [10] [11].Figure 3a) Lattice structures proposed by Steeves showing high and low CTE constituents [10]. b) Topologically optimised cell by Sigmund showing high CTE constituents in green and low CTE constituents in blue [11].The lattice structure used by Sigmund [11] was successfully printed using DED in nickel and chromium [5]. This can also be achieved with a polymer. Qang et. al. has made progress into 3D printing polyethylene glycol diacrylate (PEGDA) lattice structures with copper powder includes. The varying percentage of copper reinforcement up to 10% lowers the CTE from 1.56×10-1(1/K) to 5.1×10-5(1/K), a decrease of a third. This is enough to create a negative CTE over a 200K temperature range. The machine used is a customised form of Stereolithography [12].

This is enough to create a negative CTE over a 200K temperature range. The machine used is a customised form of Stereolithography [12]. DED using multiple materials is the next step to achieving a 3D structure with three- dimensional CTE control.Development of FEA ModelsAn auxetic solution has been adapted for maximum thermal expansion change whilst maintaining stiffness. It can be seen that the expansion rate of the outer ring is limited; however the lattice structure is allowed to expand freely. This ensures that the adjoining material (labelled L2 in figure 5a) can expand freely and bend the two neighbouring members. Both designs ensure that the angle on both lattices is equal to a tangential line, and thus maximum force is transmitted into the neighbouring structure. This is shown in figure 5a. During angle change, it is ensured that the radius of the inner annulus at Probe 1 is fixed, and the members change in length accordingly.

This is shown in figure 5a. During angle change, it is ensured that the radius of the inner annulus at Probe 1 is fixed, and the members change in length accordingly. Fixing this2141radius is required for a specific lens diameter. When the angle θ changes, the diameter of the circle attached to the tangent line also increases. These probe locations are where the displacement will be measured from in the y-direction using a Static Thermo-Structural Finite Element Simulation in ANSYS. The displacement outputs can be used to calculate the CTE and stiffness using the equations below;𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷 𝑓𝑓𝑓𝑓𝑓𝑓𝐷𝐷 𝐷𝐷𝑓𝑓𝑓𝑓𝑝𝑝𝐷𝐷 1𝐹𝐹𝑓𝑓𝑓𝑓𝐷𝐷𝐷𝐷 𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝑅𝑅 𝐷𝐷𝐷𝐷 𝐹𝐹𝐹𝐹𝐹𝐹 𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷 𝑓𝑓𝑓𝑓𝑓𝑓𝐷𝐷 𝑃𝑃𝑓𝑓𝑓𝑓𝑝𝑝𝐷𝐷 1𝐶𝐶𝐶𝐶𝐶𝐶 =𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆 =(𝑅𝑅𝐷𝐷𝑅𝑅𝐷𝐷𝑅𝑅𝐷𝐷 𝐷𝐷𝐷𝐷 𝐷𝐷𝑓𝑓𝑓𝑓𝑝𝑝𝐷𝐷 1) 𝑋𝑋 (𝐶𝐶ℎ𝐷𝐷𝐷𝐷𝑎𝑎𝐷𝐷 𝐷𝐷𝐷𝐷 𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝐷𝑓𝑓𝐷𝐷𝐷𝐷𝑅𝑅𝑓𝑓𝐷𝐷)Figure 6 provides a visual representation of the angle change from 50 to 75 degrees.

This is the minimum and maximum angle that can be used, and its effect on the CTE will be analysed. The materials used are INVAR with a CTE of 1x10-6(1/K), and Ti64Al with a CTE of 9.4x10-6(1/K). In a separate simulation where the temperature is static, loads have also been placed normal to each structure around the annulus to determine its stiffness in the elastic region. The results have been plotted on a graph shown in figures 7 and 8.Figure 4a) Shows both equal angles theta on its most efficient tangential line. b) Figure showing ANSYS deformation probe locations used to take CTE and stiffness measurements.Figure 5) Shows the structure as it changes from the minimum to maximum angle.2142Figure 6) Graph showing simulation results of the CTE range as the angle (θ) of the lattice structure changes.

Plots made for each member thicknesses (L, La).Figure 7) Graph showing simulation results of the stiffness in N/m as the lattice angle (θ) changes for each member thickness (L).By analysing the simulation data it can be seen a negative CTE can be obtained using the lattice structure. In all member thicknesses, the CTE of the bi-metallic ring at probe 2 can be seen as relatively constant. The displacement at probe 2 is steady despite the angle change since the ring cannot move; only the lattice can move inwards due to its geometry. This can be seen in figure 2b). The difference between probe 1 and 2 denotes the effectiveness of the lattice structure. It can be seen that the increase in member thickness has not seen a dramatic change in CTE, but exhibits a far higher stiffness. This is a unique feature of the structure, since the adjoining material (La) is twice the length of the member thickness.

This is a unique feature of the structure, since the adjoining material (La) is twice the length of the member thickness. The simulation results show that the relationship between the angle increase and the CTE increase produces and more stable rate of change between each angle after 60°. This is because as we reduce the angle below 60° the movement moves away from being stretch dominated to bend dominated, and the displacements that would otherwise be created and translated into stresses inside the members.

The CTE for the simulation has been taken from our testing shown in table 2. The specific stiffness is also called the stiffness to weight ratio, where weight is divided by its corresponding stiffness. This provides us with an idea of the effectiveness of the solution at keeping a lightweight design whilst remaining stiffness. The difference CTE between values at probe 1 and probe 2 provide a measure of effectiveness for the lattice structure. A higher difference implies more work done by the structure to reduce the CTE. Typically, this will be most effective where a high CTE material is present as the primary material. Aluminium coupled with Titanium or Steel as a secondary material is a good example, reducing the CTE of a conventional solid ring from 24x10-6(1/K) to -1.5x10-6(1/K) when L=1mm and θ=60°.

Aluminium coupled with Titanium or Steel as a secondary material is a good example, reducing the CTE of a conventional solid ring from 24x10-6(1/K) to -1.5x10-6(1/K) when L=1mm and θ=60°. A good choice of material is one where the combination provides the required CTE in between the range of angle, or range of thickness depending on stiffness required.Materials and ExperimentTo verify designs created through simulation data, the structures need to be manufactured and tested for performance. To ensure the most accurate simulations, the CTE of LPBF parts have been measured. This data has been fed back into the simulation as material data. It is important to measure the CTE of parts manufactured using LPBF because the CTE will differ from that of the corresponding material manufactured through traditional means. Certain parameters such as build direction and porosity have all been shown to affect the thermal expansion of a component [13][14][7].

Certain parameters such as build direction and porosity have all been shown to affect the thermal expansion of a component [13][14][7]. The CTE of the entire lattice structure also needs to be considered.Dilatometer MeasurementsTo successfully simulate and manufacture a structure capable of changing the CTE, first an understanding of the CTE of printed parts on a LPBF machines is required. Two LBPF machines are available at Swansea University, a Renishaw AM250 which utilises a 200W 1070nm Yb fibre optic laser with a laser scanning speed of up to 2m/s, and a RenAM500 which utilises a 500W laser with a similar bed size and scanning characteristics. Using the AM250 several dilatometer roads (6mm diameter x 25mm length) were built in vertical and horizontal orientations. The rods were tested using a NETZSCH 402 dilatometer featuring a2144Weight (g)92.963.75127.5273.454.2436.7567.850.32cryogenic furnace capable a temperature range of -200°C and +1000°C.

These results are given in table 2 for a temperature range of -70°C and +100°C from a sample size of -100°C to +130°C. Trimming 30°C off both extremities removes noise from the data set.Figure 8) Graphs showing the CTE against temperature built vertically and horizontally. Measured using the NETZSCH 402 Dilatometer.Material Titanium Ti64Al INVAR Aluminium [15]Temperature Range -70°C to +100°C -70°C to +100°C 25°C to 200°CH CTE 9.12x10-6(1/K) 1.31x10-6(1/K) N/AV CTE 9.00x10-6(1/K) 1.50x10-6(1/K) N/AAveraged CTE 9.06x10-6(1/K) 1.41x10-6(1/K) 24x10-6(1/K)Table 2) Shows dilatometer results for INVAR and Ti64Al using the Renishaw AM250, and Renishaw RenAM500 LPBF Machine respectively. Aluminium is referenced from another source.This data is very important to determining the most effective method of controlling CTE. It can be seen that the CTE for INVAR increases as the temperature decreases.

It can be seen that the CTE for INVAR increases as the temperature decreases. This is different to the CTE for Ti64Al which decreases as the temperature decreases. This will need to be taken into account with any CTE offsetting design. Heat treatment of the part can change the thermal expansion values, but for this experiment we are looking at as built.Results and DiscussionIt is observed that an optimum lattice angle for the proposed design is between 60-75degrees. It is observed that the stiffness is increased greatly as the member thickness (L) is increased, but the CTE only decreases marginally. This is due to the cross-member lengths La being twice that of L. The force created from the expansion of adjoining members (La) must overcome equally the additional thickness of members (L). Optical engineers will be able to choose an angle, thickness (L), and material combination from experimentation following the design rules proposed.

Optical engineers will be able to choose an angle, thickness (L), and material combination from experimentation following the design rules proposed. It can be seen that the high specific stiffness of 5.2.E4.m2 s-2 for aluminium made it a great choice for lightweight applications. The combination of aluminium and titanium in the proposed lattice structure features one of the highest reductions in CTE from 24×10-6(1/K) to -1.5×10-6(1/K).Conclusion2145Modelling shows that an auxetic lattice of a single material coupled with an external containing ring can be used to create customized coefficients of thermal expansion around a 2D annulus. This opens the possibility of matching CTE’s within an optical system to the lens material. It has shown that a design engineer can follow a set of design rules regarding member thickness and lattice angle to achieve the desired co-efficient of thermal expansion.

It has shown that a design engineer can follow a set of design rules regarding member thickness and lattice angle to achieve the desired co-efficient of thermal expansion. Negative CTE’s have been achieved when using a high CTE ratio such as INVAR and titanium. Further work is required to test the designs and ensure they match the simulation data. The future multi-material print requires further work in manufacturing with Fused Deposition Modelling (FDM) polymer machines featuring different amounts of carbon and glass fibres. A DED LENS system will also be explored for making a multi-metallic print.AcknowledgementsThis research was supported by Sêr Cymru funding through the National Research Network in Advanced Engineering and Materials in Wales (grant NRN145) and by Qioptiq.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperConvection Heat Transfer Coefficients for Laser Powder Bed FusionMohammad Masoomi, Arash Soltani-Tehrani, Nima Shamsaei, Scott M. Thompson*Department of Mechanical Engineering, Auburn University, Auburn, AL 36849National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 3684929th Annual International Solid Freeform Fabrication Symposium - An Additive Manufacturing ConferenceAugust 2018AbstractThis study investigates the effects of convection heat transfer during the laser-powder bed fusion (L-PBF) additive manufacturing process. For the L-PBF process, parts are fabricated under a purged, inert environment to avoid oxidation. Part of the delivered laser energy is transferred to the process chamber/environment through radiation and convection during fabrication.

Part of the delivered laser energy is transferred to the process chamber/environment through radiation and convection during fabrication. In this study, customized computational fluid dynamics (CFD) software is used to simulate the L-PBF of a single layer of stainless steel 17-4 PH. Local temperature and temperature gradients, as well as dimensionless numbers descriptive of important thermophysics, are provided in order to quantify local convective heat transfer. The results are used to predict local heat transfer coefficients during the L-PBF additive manufacturing process.KEYWORDS: Laser powder bed fusion, Convective heat transfer coefficient, Temperature gradient, Numerical simulation1.IntroductionInert shielding gas is employed during welding and laser-based additive manufacturing (AM) processes to reduce oxide formation along part/powder surfaces, protect the melt pool against contamination and to aid in expelling so-called ‘spatter’ debris.

Effects of shielding gas flow on welding/AM processes and on final material quality have been studied. Ly et al. [1] studied spatter generation during laser-powder bed fusion (L-PBF) process via experimentation and it was shown that the presence of flowing argon gas can decrease spatter significantly relative to when the process is performed in vacuum. Wang et al. [2] studied the effects of shielding gas during laser deep penetration welding. It was shown that the shielding gas can help stabilize the laser welding process via convection heat transfer and enhanced ionization–recombination. Philo et al. [3] simulated the flow within an L-PBF (Renishaw AM 250) chamber. The flow atop the powderthecorresponding author: smthompson@auburn.edu1686bed was shown to be non-uniform and these flow non-uniformities increased spatter generation. Kah et al.

The flow atop the powderthecorresponding author: smthompson@auburn.edu1686bed was shown to be non-uniform and these flow non-uniformities increased spatter generation. Kah et al. [4] experimentally studied the effects of various shielding gases on the microstructure and mechanical properties of different materials including steel and aluminum during welding. For austenitic stainless steels, it was found that increasing the amount of nitrogen in the shielding gas can increase the ductility and improve the tensile strength, hardness, and pitting corrosion resistance of the final weld [4].The L-PBF process may be numerically modeled and simulated using finite element (FE) methods [5]–[8]. Such approaches have been performed with varying degrees of fidelity for learning process-structure-property relationships while minimizing the need for repetitious, costly experimentation.

Such approaches have been performed with varying degrees of fidelity for learning process-structure-property relationships while minimizing the need for repetitious, costly experimentation. In general, a continuum or powder-scale approach can be utilized, with the latter requiring significantly more computational resources. In order to produce realistic boundary conditions for performing more accurate laser-based AM simulations, the convection and radiation heat transfer along the surfaces of as-manufactured parts and surrounding powder should be appropriately considered.In order to better represent boundary conditions, effects of convection need to be taken into account. Several different techniques and approaches have been used. In some cases, radiation and convection have been assumed as negligible during L-PBF [9]–[11], while others have assigned a uniform/constant heat transfer coefficient over all free surfaces. Peyre et al.

Peyre et al. [12] used a constant heat transfer coefficient of 20 W/m2∙ºC, Dong et al. [13] used 25 W/m2∙ºC, Espana et al. [14] used 10 W/m2∙ºC, while Denlinger et al. [15] and Parry et al. [16] used 15 W/m2∙ºC. Michaleris [7] assumed free convection around the powder bed and forced convection around the heat affected zone (HAZ), i.e. 10 W/m2∙ºC and 21 W/m2∙ºC, respectively. It was concluded that neglecting the effects of convection and radiation results in an over-prediction of temperature. More details on convection heat transfer modeling during L-PBF have been presented by Masoomi et al. [17].The current study focuses on modeling the thermal response of a single track during typical L-PBF conditions while considering convection heat transfer with the inert shielding gas. The gas momentum and energy fields are found numerically while referencing temperature-dependent properties.

The gas momentum and energy fields are found numerically while referencing temperature-dependent properties. Effects of convection and radiation heat transfer on the temperature distribution and temperature gradient of SS 17-4 PH during L-PBF are investigated. The difference between simulating L-PBF with an assumed constant heat transfer coefficient versus locally-variable heat transfer coefficients, found via CFD of surrounding fluid flow in the chamber, are also elucidated. Finally, data from simulations were used to propose a local heat transfer coefficient for the HAZ and rest of the build plate and results are compared to other simulations. In this study, effects of evaporation in the melt pool thermal response and its effects on convection have been neglected. Nonetheless, this study is believed to be a critical first step to understanding convection heat transfer due to shielding gas during L-PBF.2.

Nonetheless, this study is believed to be a critical first step to understanding convection heat transfer due to shielding gas during L-PBF.2. Physical model and numerical methodThe transient heat transfer during L-PBF of metals is complex; exhibiting dense heat flux transport, inhomogeneous phase change, small length/time scales, melt pool instability,1687surface/fluid interaction, microstructural coupling, anisotropic microstructural growth and more [18], [19]. Utilizing a physical model that reduces this complexity can be advantageous for minimizing computational investment and for providing general insight into the process-property relationships inherent to parts fabricated via L-PBF. One such approach is to model the heat in all participating media while neglecting, or creatively accounting for, diffusion secondary/tertiary effects such as melt pool flow dynamics, powder bed mechanics, powder heterogeneity and microstructural evolution (i.e. solid-phase transformations).

solid-phase transformations). The melt pool, powder bed and solidified part can then be more readily modeled as continua with effective density and transport properties, and this approach is adopted herein. Melt pool advection and phase- change are accounted for by employing an effective melt pool thermal conductivity and in estimating latent heat transfer by computing solid/liquid phase fractions. Solid-phase nucleation, liquid-solid wetting phenomena, pore generation and the effect of powder size distribution are neglected. Effects of convection were modeled based on the technique presented by Masoomi et al. [17]. The shielding gas type was assumed to be argon with temperature-dependent properties.3. Results and DiscussionTetrahedral meshes were used for the powder bed and were uniformly dispersed at a size of ~3800 μm3 (mesh lengths of ~35 μm). This mesh size was found to provide for mesh-independent results. The time step was chosen as 1 ms.

This mesh size was found to provide for mesh-independent results. The time step was chosen as 1 ms. Process parameters for fabrication of SS 17-4 PH are presented in Table 1. For all simulations the argon gas and laser were moving in the same direction.Table 1. Process parameters for 17-4 PH stainless steel provided by EOS.Laser power (W) Scan speed (mm/s) Hatching distance (μm) 220755.5100Layer thickness (μm)40Figure 1 presents the local temperature and temperature gradient response at the track’s mid- point when the gas inside the chamber is modeled, Fig 1 (a), and when a constant heat transfer coefficient of 20 W/m2∙ºC, is used. It may be seen that the maximum temperature of the melt pool drops around 70 ºC (~ 3%) when the convection is calculated directly by modeling gas inside the chamber. This is due to a better prediction of convection and more accurate description of boundary conditions. This same trend is observed with the local, absolute temperature gradient.

This is due to a better prediction of convection and more accurate description of boundary conditions. This same trend is observed with the local, absolute temperature gradient. The temperature gradient is approximately 20% lower when the gas inside the chamber is modeled compared to the case where a constant heat transfer coefficient is used.1688Figure 1. Temperature and temperature gradient at the middle of track when (a) convection is directly modelled in the chamber, (b) a constant heat transfer coefficient (h = 20 W/m2∙ºC) is used.The Nusselt number is used to aid the characterization of convection heat transfer duringthe fabrication of parts, i.e. Eq. 1:ℎx𝑥𝑥 𝑘𝑘gNux =where hx is the local heat transfer coefficient and kg is the thermal conductivity of the shielding gas. The local heat transfer coefficient was calculated using:h𝑥𝑥 = −𝑘𝑘g 𝜕𝜕𝜕𝜕 𝜕𝜕𝜕𝜕 �𝜕𝜕s,x − 𝜕𝜕∞��𝑧𝑧=0where Ts,x is the local surface temperature.As shown in Fig.

The local heat transfer coefficient was calculated using:h𝑥𝑥 = −𝑘𝑘g 𝜕𝜕𝜕𝜕 𝜕𝜕𝜕𝜕 �𝜕𝜕s,x − 𝜕𝜕∞��𝑧𝑧=0where Ts,x is the local surface temperature.As shown in Fig. 2, the Nusselt number is higher near the HAZ during L-PBF. Based on the plot, and by using Eq. (2), the average heat transfer coefficient near the HAZ may be found to be ~1000 W/m2∙ºC and for the rest of surface area at ~50 W/m2∙ºC.1689(1)(2)Figure 2. Nusselt number at the middle of track when convection is directly modelled in the chamberSimulations were repeated by employing new heat transfer coefficients for the HAZ, at 1000 W/m2∙ºC, and rest of the area, at 50 W/m2∙ºC. The temperature and temperature gradient when using these modified heat transfer coefficients are presented in Fig. 3. It may be seen that by using the modified heat transfer coefficient, the accuracy for peak temperature increases by 2%.

3. It may be seen that by using the modified heat transfer coefficient, the accuracy for peak temperature increases by 2%. It may be seen that the maximum temperature changes by less than 20 ºC when using the modified heat transfer coefficients. Additionally, the temperature gradient shows 10% improvement relative to using only a constant heat transfer coefficients. It is important to note that modeling the gas flow inside the chamber is computationally expensive and using modified heat transfer coefficients is more efficient.Figure 3. Temperature and temperature gradient at the middle of track when convection is directly modelled in the chamber16904. ConclusionsA continuum-based, numerical model was used to study the effects of shielding gas convection on total heat transfer during the L-PBF of SS 17-4 PH. The convective heat flux was calculated directly by modeling typical gas flow characteristics within the L-PBF chamber via computational fluid dynamics (CFD).

The convective heat flux was calculated directly by modeling typical gas flow characteristics within the L-PBF chamber via computational fluid dynamics (CFD). The gas velocity and temperature boundary layers atop the powder bed and track were calculated. The combined melt pool/HAZ temperature distribution was found while either: (i) considering the gas-flow-coupled convective heat flux, (ii) constant heat transfer coefficient of h = 20 W/m2·ºC, or (iii) considering a modified heat transfer coefficient of h = 1000 W/m2·ºC for HAZ and h = 50 W/m2·ºC elsewhere. The major findings from the study are summarized below:1) Modeling the gas inside the chamber will result in a more accurate prediction of thermal response during the L-PBF process compared to when using a constant heat transfer coefficient.2) Convective heat transfer is more prominent near the HAZ compared to the rest of build area.

3) Using a modified heat transfer coefficient will improve the accuracy of thermal simulation compared to the case when a constant heat transfer coefficient is used.4) The temperature gradient is more sensitive to convective heat transfer compared to maximum melt pool temperature.Further studies can be done to better determine the effects of flow direction, gas speed and type of gas on convective heat transfer coefficients during L-PBF. A formula can be used to predict revised heat transfer coefficients based on the environmental conditions.5. AcknowledgementsThis work was partially supported by the National Science Foundation under Grant #1657195, and the National Aeronautics Space Administration (NASA) Grant #80MSFC17M0023.References [1]S. Ly, A. M. Rubenchik, S. A. Khairallah, G. Guss, and M. J. Matthews, “Metal vapor micro-jet controls material redistribution in laser powder bed fusion additive manufacturing,” Sci. Rep., vol. 7, no. 1, p. 4085, Dec. 2017.[2] H.

nverting a CAD model into a manufacturing model for the components made of a multiphase perfect materialFeng Zhua, Ke-Zhang Chena, Xin-An Fengb aDepartment of Mechanical Engineering, The University of Hong Kong, Hong Kong bFaculty of Mechanical Engineering, Dalian University of Technology, China Reviewed, accepted August 4, 2004AbstractTo manufacture the component made of a multiphase perfect material (including homogeneous and multi heterogeneous materials), it CAD model should be processed and converted into layered manufacturing model for further transformation of numerical control (NC) coding. This paper develops its detailed approaches and corresponding software. The process planning is made first and includes: (1) determining the build orientation of the component; and (2) slicing the component into layers adaptively according to different material regions since different materials have different optimal layer thickness for manufacturing.

After the process planning, the layered manufacturing models with necessary information, including fabrication sequence and material information of each layer, are fully generated.1. IntroductionWith the rapid development of high technology in various fields, there appear more special functions of the mechanical components or products, which may require component materials to possess some special properties. Since conventional homogeneous materials cannot satisfy these requirements, attention has been paid to heterogeneous materials, which may include composite materials, functionally graded materials (FGMs) and heterogeneous materials with a periodic microstructure (HMPMs), which are shown in Figure 1. However, different portions of a component may have different special requirements and the component made of a single heterogeneous material cannot meet all special requirements in its different portions, but only some.

To satisfy all requirements, it would be necessary to use components made of different materials, including homogenous materials and the three types of heterogeneous materials, thus meeting all special requirements in different portions and also making the best use of different materials, just like nature’s organisms (e.g. bamboo, tooth, bone etc.) which have perfect combination between materials and functions after a long time of evolution. A component, which has a perfect combination of different materials (including homogeneous materials and different types of heterogeneous materials) in its different portions for a specific application, is considered as the component made of a multiphase perfect material (CMPM).(a)532(c)(b)Fig.

1 Three types of heterogeneous materials To design and represent such components according to the requirements from high-tech applications, a corresponding computer-aided design method [1] (including both geometric and material design) and a corresponding CAD modeling method [2, 3] (containing both geometric and material information) have been successfully developed and can be implemented by applying the functions of current CAD/CAE software.2.

Review of the fabrication approach for the CMPMsTo produce the CMPM, the fabrication technique should satisfy the following requirements [4]: (1) Adding materials that may have different constituents with their required volume fractions for each pixel (needed for all these heterogeneous materials) and inclusions with the required distribution and quantity (needed for composite materials); (2) Obtaining the material layer (probably very thin) with precise thickness (needed for periodic microstructures) and boundary(needed for all these heterogeneous materials); (3) Removing materials from the layer to create very small and precise voids for periodic microstructures (needed for the materials with a periodic microstructure).It is obvious that the CMPM cannot be fabricated using traditional material subtractive processes (e.g. machining) or material additive processes (e.g. layered manufacturing (LM)).

machining) or material additive processes (e.g. layered manufacturing (LM)). A hybrid manufacturing technique incorporating with traditional machining, LM and micro- fabrication techniques was proposed [4].

The workflow of this hybrid LM technique is as follows:(1) If there is adjoining material regions which are higher than the layer to be spread, remove the superfluous material from the layer by an end mill to obtain precise boundary of the layer, since the spraying area of a jet is much larger than a pixel so that the practical area of obtained layer is larger than the required area, and, at the same time, suck out the formed chips by vacuum;(2) Spread a material layer with the required thickness (layer thickness must plus a rough grinding allowance and a finish grinding allowance for the material region of periodic microstructures) and constituent composition (for all material regions) and, at the mean time, spray the inclusion with the require distribution and quantity to stick in the layer (only for composite material region) for every pixel;(3) Grind the top surface of added material layer by the annular end face of a cup grinding wheel to obtain a required thickness of material layer (layer thickness must plus a finish grinding allowance for the material region of periodic microstructures), since metal cladding is not flat enough and the thickness of the added layer is not accurate enough, and, at the mean time, suck out the formed chips or sludge by vacuum; If the layer being fabricated is not the material region of periodic microstructures go to step (1).(4) Engrave or sculpt the layer to create the necessary voids for periodic microstructures with a required depth (layer thickness plus a finish grinding allowance), and, at the mean time, suck out the formed chips or sludge by vacuum.(5) Fill the void with a material with lower strength and high melting point to avoid the refillingof the material spray for next layer;(6) Grind the top surface of the material layer again to remove superfluous lower strength material and the burrs formed (3).

This will ensure that the required thickness and flatness of the material layer is produced;533(7) Repeat step (1) to spread material for next layer until the component is completely fabricated. The component so formed will have the required constituent composition, inclusions and their distribution, and/or periodic microstructures with lower-strength materials in the voids. The lower-strength metal in the voids will not affect the function of the component and can protect the component from erosion.In the second step, plasma spraying [5] has been selected for the spreading and Lasermicromachining [6] is used for the engraving in the fifth step.3. Process planning for the CMPMsBased on the hybrid fabrication technique, the process planning (PP) for the CMPMs can be further developed. The PP acts as a bridge between the 3D CAD model domain and 2.5D layered manufacturing model domain.

The PP acts as a bridge between the 3D CAD model domain and 2.5D layered manufacturing model domain. Although the CMPMs cannot be produced by applying the current LM machines, they are still fabricated layer by layer in the hybrid manufacturing process. Therefore, the PP for the CMPMs can be developed based on that of the traditional LM. The PP for the traditional LM goes through: (1) determining the build orientation of an object; (2) slicing the object; (3) generating the support structure for the object; and (4) generating the tool paths. After the first two steps, the 3D model has been converted into 2.5D layers, and the necessary information has been generated. Since the last two steps can be performed based on the generated layers and usually implemented by commercial software packages, the last two steps are not discussed in this paper.Many algorithms for the PP in LM have been reported in some literatures [7, 8].

However, these methods are mainly based on an object’s geometry and do not consider the material variations, inclusions and microstructures in the heterogeneous objects. Some researchers [9] proposed a new approach for objects with FGM by taking into account the material resolutions and the varying direction of material volume fractions. But the machinability and the effective properties of microstructures still have not been considered. Thus, a new approach for CMPMs should be developed, in which the first priority should be to ensure its functional properties, i.e. material effective properties, and the geometric accuracy can be regarded as the second priority.3.1 Determining the build orientation of an objectThe purpose of the first step is to select a suitable direction in which the object is to be fabricated.

During determining this direction, many related factors need to be considered, such as the total build time, surface quality and finish, component mechanical properties etc [8]. The orientation of the objects made of FGM in some powder-based LM processes is determined mainly based on reducing build time and material constituent composition error. In these processes, a heterogeneous object is discretized into a lot of small homogeneous layers, each of which has a specified mixed material without gradation. All of the layers are deposited one by one and the powder for the specified mixed material of each layer is fed manually. Therefore, the goal is met by selecting a suitable direction that can minimize the variation of material constituent composition in each layer [9], which is also the best way to ensure the effective properties of the finished FGM [10].

But, using the plasma spraying in the hybrid manufacturing system, the material powders are mixed automatically according to the requirements and deposited pixel by pixel. Therefore, the objects’ quality and build time are not effected by the534varying direction of material constituent composition.In the case of CMPMs, the machinability and the effective properties of periodic microstructures have to be considered further. This requirement can be satisfied by minimizing following summation (∑∫ αi iα is the angle between surface Ai and the normal of the worktable of LM machine. The approach can increase layer’s thickness in LM machine, thus can improve the fabrication efficiency.iA ), where Ai is the area of the main surfaces in base cell, and3.2 Adaptive slicingAfter the build direction has been determined, the orientated model is sliced to yield the slices.

To slice the objects with homogeneous materials, only geometric surface finish of the object and build time are considered. The contradiction between these two factors led to the development of different slicing algorithms [12]. However, these algorithms cannot be directly applicable to heterogeneous objects, and have to be modified to take into account the material information [13]. Kumar et al. [13] proposed an approach to address the issues of slicing heterogeneous objects by considering both the geometric profiles and material variations. But, for object with multi heterogeneous materials, the reasonable layer thickness for periodic microstructures and composites should also be considered, respectively, because different heterogeneous materials have different optimal layer thicknesses for manufacturing. Figure 2 illustrates different layer thickness for different material regions (denoted with different colors and symbols).

Figure 2 illustrates different layer thickness for different material regions (denoted with different colors and symbols). It should be emphasized that a “layer” in this paper refers a material slice for one type of material region, not for the whole component. Therefore, a new slicing algorithm needs to be developed. Before the slicing operation, the component should first be divided into blocks using parallel planes perpendicular to the build orientation to ensure that each block only has the (part of) material regions whose height equal the height of block. Since the slicing operation is associated with fabrication sequence determination, its detail will be discussed in the next section.Fig.2 Different layer thickness for different material region4. Establishing the layered manufacturing modelAlthough the CAD model for CMPMs include all the necessary information, it cannot be applied for layered manufacturing directly.

Establishing the layered manufacturing modelAlthough the CAD model for CMPMs include all the necessary information, it cannot be applied for layered manufacturing directly. The 2.5D layers with geometric and material information should be integrated as a layered manufacturing model, so that the numerical control codes can be programmed. The layered manufacturing model has three types of data: (1) Fabrication sequence; (2) Spraying information; and (3) Periodic microstructures information. The approaches for generating these data are introduced as follows:5354.1 Fabrication sequenceAs mentioned previously, for the object with multiphase perfect materials, the reasonable layer thickness for different material regions may be different. In other words, the spraying surface of different material regions are usually not identical. The spraying thickness of periodic microstructures regions should be determined by the size and complexity of their base cells.

The spraying thickness of periodic microstructures regions should be determined by the size and complexity of their base cells. The base cells with smaller size and more complicated shapes have smaller thickness. This thickness can be assigned as 1/10 ~ 1/20 of the base cell dimension, and usually less than 0.1 mm. The spraying thickness of composite materials and FGM regions should be determined based on the variations of material constituent compositions, and typically 0.5 ~ 2mm. The region with larger gradient of material constituent compositions has smaller thickness. Since the top surface of added material will be grinded by an annular face of a cup grinding wheel after the spray operation, the material surface to be grinded should be higher than others. Therefore, a reasonable fabrication sequence planning is imperative.

Therefore, a reasonable fabrication sequence planning is imperative. The procedure goes through (1) Sub-dividing the CAD model into blocks; (2) Generating fabrication sequence for each block; and (3) slicing the blocks into layers, which is introduced in detail as follows:4.1.1 Sub-dividing the CAD model into blocksIn order to calculate the deposition thickness and plan the fabrication sequence, the object should be first divided into several blocks by parallel planes perpendicular to the build orientation. Each block only has (part of) the material regions whose height is equal to the height of block. This can be done using a sub-division algorithm. For example, the CAD model for a CMPM shown in Figure 3 has five material regions and can be divided into four blocks by using this algorithm.Fig.

This can be done using a sub-division algorithm. For example, the CAD model for a CMPM shown in Figure 3 has five material regions and can be divided into four blocks by using this algorithm.Fig. 3 Sub-dividing the CAD model into blocks4.1.2 Fabrication sequence generation for each blockSince each block only has such material regions whose heights are equal to the height of block, the deposition thickness and fabrication sequence for one work cycle in each block can be calculated. For each block, the input data are its CAD model and layer thickness for each material region in this block. The output data are the deposition thickness and the fabrication sequence in each work cycle. For example, assume that the optimal spraying thicknesses for periodic microstructures, FGM and composite materials are 0.2mm, 0.5mm and 1.5mm, respectively. Their lease common multiple is 3.0mm, which can be considered as the deposition thickness for a work cycle.

Their lease common multiple is 3.0mm, which can be considered as the deposition thickness for a work cycle. The fabrication in one work cycle is illustrated in Figure 4. The cycle has 18 steps,536and in each step the worktable moves down 0.1 or 0.2 mm. Once the worktable moves down 0.2mm, the periodic microstructure regions denoted by M will be sprayed. When the worktable moves down every 0.5mm, the FGM regions (denoted by F) will be sprayed. And when the worktable moves down every 1.5mm, the composite material regions (expressed using C) will be sprayed. In the case that different material regions have the same spraying height, the spraying operations are executed for different material regions one by one. In Figure 4, the hatches denote the spraying thickness.M F CMMMFM)m mM F(MFMMM FMMF CMMe c n a t s i d g n i d n e c s e dMspread layersequence numberFig 4.

Fabrication sequence in one work cycle4.1.3 Fabrication sequence generation for the objectHaving used above mentioned approach to generate the fabrication sequence for each block, the fabrication sequence for the object can be generated by integrating the fabrication sequence of all blocks.4.1.4 Adaptive slicing operationsFollowing the generated fabrication sequence, the adaptive slicing routine is called to yield the geometric profiles, generate spraying information and microstructures information for each material region at the calculated height, which have been reported in the literature [14], and outlined in the following sections.4.2 Spraying informationThe spraying data on each layer includes geometric and material information: i.e. the geometric boundaries of the spraying regions; the spraying thickness of each spraying region; the code names of material constituents and their compositions; the code names of composite inclusions and their volume fractions.

These two types of information are described as follows:4.2.1 The geometric boundaries of material regionsAs discussed in adaptive slicing, the geometric boundaries can be obtained by intersecting the3D geometric models of each material region in the CMPMs with the slicing plane.5374.2.2 The material constituents and their volume fractions in each spraying regionThe code names of material constituents and their compositions and the code names of composite inclusions and their volume fractions can be retrieved from the CAD model of the CMPM. The constituent composition function in the CAD model, which is the function as to the coordinate values of spraying points (x, y, z) in the local coordinate system, should be first converted into the function as to the points (X, Y, Z) in the global coordinate system.

Since the CAD model intersects with the plane Z=Z0 (Z0 is the height of the spraying surface), the final constituent composition function has been converted to a function as to 2D points (X, Y) in the global coordinate system.4.3 Periodic microstructures informationThe periodic microstructures information in each layer includes: spraying thickness, the 2D shape and position of the voids. The periodic microstructures information in the CAD model for the CMPMs covers the variational geometric model of the base cell, inserting functions of insertion positions, functions for calculating dimension and orientation, with which the entire region of the periodic microstructures can be described.

The shape and position of the voids can be obtained by intersecting the base cell array of microstructures, which are determined by their inserting points, with the spraying plane.Since the number of inserting points of a material region is very huge while the number of base cells intersected with the spraying plane is much less comparatively, an efficient scanning algorithm is required to reduce the time of searching the related inserting points. This paper applies an approach [14], with which the search space is restricted within a very narrow scope for saving computing time. As shown in Figure 5, the inserting point array is sliced by a spraying plane with a thickness S and an angle Φ from positive z-axis. Two auxiliary planes parallel to the spread or slicing layer are used to determine the search scope.

Two auxiliary planes parallel to the spread or slicing layer are used to determine the search scope. The distance between each auxiliary plane and the middle layer of spread layer is Lmax, which is the largest difference of the vertical distances of inserting points from the middle layer of spread layer. This can ensure that all the base cells to be sliced on the layer should fall into the space between the two auxiliary planes.zLmaxupper auxiliary planeLmaxlayer to be spreadmiddle layersLmaxlower auxiliary planeFig 5. Search the base cells intersecting with the spraying plane5385. Data structure of layered manufacturing modelAfter generating the layered manufacturing model, the geometric and material information for each layer should be integrated and kept in the layer files, with which the NC codes can be programmed. Currently, many layer formats are available, including CLI, LEAF, SLC, HP-GL, and G_WORP etc. [15].

Currently, many layer formats are available, including CLI, LEAF, SLC, HP-GL, and G_WORP etc. [15]. However, they only carry the geometric information, and hence cannot be applied directly as the layer format for the layered manufacturing model of CMPMs. As to those objects with material information, there is no suitable layer format to represent them yet, and this issue has actually become one of the bottlenecks in the rapid prototyping industry [7]. Aiming at this point, a frame of new data structure for the layered manufacturing model of CMPMs is proposed as shown in Figure 6. The information can be organized hierarchically into a tree structure.CAD modelBlock 1Block 2…Block nWork cycle 1_1Work cycle 1_2Work cycle 1_n…Work cycle n_1…Work cycle n_nLayer 1_1_1Layer 1_1_2…Layer 1_1_n…Layer n_n_1 …Fig. 6. The tree structure of the CAD modelIn the model tree, each leaf is a layer, which is kept in a file.

6. The tree structure of the CAD modelIn the model tree, each leaf is a layer, which is kept in a file. Note that not all of the layers should be stored, since the data of the whole CAD model is very huge. A “use and throw” strategy can be employed, in which only one layer is calculated and stored at a time. That is, when the current layer is under fabrication, the next layer is calculated and stored. When the fabrication moves to the next layer, the data for the current layer will be discarded and another layer will be calculated and stored, until all of the layers are finished.A layer is the smallest unit in this data structure, which can be regarded as a class (shown inFigure 7) according to the object-oriented thought [16].Layer AttributesLayer attributesGeometric attributesMaterial attributesServicesCall related functions539Fig. 7 A layer classLayer n_n_nFor example, as to the FGM layer, its class diagram is illustrated in Figure 8.

7 A layer classLayer n_n_nFor example, as to the FGM layer, its class diagram is illustrated in Figure 8. When a material constitute composition function is specified, its corresponding sub-routine is called, the boundaries and material code names of one layer are inputted as parameters, then the material volume fractions at each position can be calculated correctly.FGM LayerAttributesCode namesLayer no.HeightThicknessBoundariesMaterial constituent composition function no.Sub- routineLayer data fileServicesCall material constituent composition functionsFig. 8 The class for a FGM layerThe class for a composite layer has a similar diagram as that of a FGM layer, as shown inFigure 9.Composite LayerAttributesLayer no.BoundariesCode name for inclusionsHeightThicknessSpraying function no.Sub- routineLayer data fileServicesCall spraying functionsFig.

9 The class for a composite layer540As to the HMPM layer, the data structure is quite different, as the layer has a lot of micro voids that cannot be specified and calculated by functions. Therefore, a bit-map like file is introduced to describe a HMPM layer. The format consists of a header file and a data file (Figure 10). In the header file, some layer information, such as layer no., thickness and height, is recorded; while in the data file, all of the material points are denoted by “1” and void points are denoted by “0”. Thus the entire layer is described using a bit-map.HMPM LayerHeaderLayer no., Height, ThicknessDataFig. 10 The HMPM layer format6. Implementation and exampleBased on the new modeling method for CMPMs [2], a CAD modeling system was developed [3].

10 The HMPM layer format6. Implementation and exampleBased on the new modeling method for CMPMs [2], a CAD modeling system was developed [3]. With the system, users can apply the functions of current CAD graphic software to build CAD models for their components designed with a multiphase perfect material and visualize both material and geometric information for any layer of the components they select.Now the system has been extended for process planning and establishing LM models tasks. Using the system, users can determine the build orientation and divide the object into several blocks. Then the fabrication sequence is generated for each block and the CAD model can be converted into the LM model by adaptive slicing. All the data in the LM model can be kept using the proposed data structure.Figure 11 shows its user interface, where the CAD model of a pipe in the work window ① is being processed and converted into a LM model.

The internal surface of the pipe is made of a HMPM while its outside part is covered with a composite material, and the middle part is designed with FGM. Users can choose the orientation and layer thickness by dragging the scroll bar on the control panel ②, with which the slicing routine is carried out. After that, users are allowed to choose a work cycle to do fabrication sequence generation. The result is displayed in the work window ③ and all the information for LM model is stored in the form of the corresponding data structure.541Fig. 11. The user interface7. ConclusionsTo manufacture the CMPMs, a hybrid manufacturing approach has been developed. Based on this approach, this paper further proposes corresponding process planning techniques, which include build orientation determination and adaptive slicing, with which, the 3D CAD model is converted into 2.5D layers.

Then the geometric and material information associated to the layers is integrated as a layered manufacturing model, consisting of three types of data: (1) fabrication sequence; (2) spraying information; and (3) periodic microstructures information. With these data, the NC codes can be programmed properly.AcknowledgementsThe reported research is supported by Competitive Earmarked Research Grant of Hong Kong Research Grants Council (RGC) under project code: HKU 7062/00E. The financial contribution is gratefully acknowledged.542Reference1. Chen K.Z. and Feng X.A. Computer-aided design method for the components made of heterogeneous materials. Computer-Aided Design, 2003, Vol.35, pp.453–66.2. Chen K.Z., Feng X.A. CAD modeling for the components made of multi heterogeneous materials and smart materials. Computer-Aided Design, 2004: 36(1): 51-63.3. Zhu F and Chen K.Z. A CAD modeling system for the components made of multi heterogeneous materials.

OPERATIVE FABRICATION METHODOLOGY FOR EMBEDDING WIREON CURVED SURFACESC.Y. Kim, A. Cuaron, M.A. Perez, D. Espalin, E. MacDonald, and R.B. WickerW.M.Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, TX79968AbstractIn conventional additive manufacturing (AM), an object is fabricated by depositing material in a layer by layer fashion. Typically, this process is retained so that deposition can occur on flat surfaces and motion can be constrained to requiring only three degrees of freedom (DOF) in a Cartesian coordinate system. When incorporating wire in three-dimensional (3D) objects, there is sometimes a need for placement along curved surfaces on which positions are defined not only by 3D Cartesian coordinates but also angular ones. Therefore, a minimum of two additional DOFs are required allowing movement to be generated at the build platform as well as of the extrusion head.

Therefore, a minimum of two additional DOFs are required allowing movement to be generated at the build platform as well as of the extrusion head. This paper addresses a method for trajectory planning of both systems, that is, the extrusion head and the movable build platform, allowing for cooperative and harmonic motion between the two.IntroductionMaterial extrusion additive manufacturing (AM), more commonly referred to as Fused Deposition Modeling (FDM) extrudes a thermoplastic through a small diameter tip and is placed on a planar build surface to build an object one layer at a time. This manufacturing technology is rapidly growing in various applications [1][2]. To increase the diversity of applications this technology can be used in, the researchers have explored the possibility of multi-material material fabrication and integration of embedded wire.

To increase the diversity of applications this technology can be used in, the researchers have explored the possibility of multi-material material fabrication and integration of embedded wire. Previous research developed a multi- functional, multi-material system with the capability to fabricate with four different materials [3]. Successive research manufactured a CubeSat subsystem containing an FDM-build substrate, commercial off-the-shelf electronic components, and conductive ink traces [4]. This same work demonstrated an alternative to direct write of conductive inks for creating traces, which involved the ultrasonic embedding of solid copper wire. Another method for creating traces is metal fused deposition modeling, and although this method shows enhanced conductivity, challenges still exist in depositing metals on 3D surfaces.

Another method for creating traces is metal fused deposition modeling, and although this method shows enhanced conductivity, challenges still exist in depositing metals on 3D surfaces. This obstacles must be resolved as the ability to embed metal on 3Dsurfaces of FDM-fabricated parts is essential in advancing research in the 3D printed electronics and mechanics field. Therefore flexible tooling FDM technologies, which can deposit a material along various angular positions, were introduced. Song et al. attempted to make a 6 DOF FDM system using a Stewart platform integrated into the tool head, allowing for flexible deposition of material [5]. Lee et al.

Song et al. attempted to make a 6 DOF FDM system using a Stewart platform integrated into the tool head, allowing for flexible deposition of material [5]. Lee et al. used a 5 axis desktop CNC, converted into an FDM machine, in which the tool head was able to move linearly in the z and x directions, while rotating along the y-axis and the base moved in the y direction while rotating along the x-axis [6].To optimize embedding metal onto the surface of an object, the tool head must be perpendicular to the surface of the model base. For that to be maintained, an actuator is required to change the orientation of each tool. Although conventional 3D printers have 3 DOFs equipped with anactuator, this is not enough to maintain the tool perpendicular. Therefore, to embed on a curved surface, a minimum of 2 additional DOFs are required. In this paper, it is suggested that 2 DOF actuators will be added to the build platform.

Therefore, to embed on a curved surface, a minimum of 2 additional DOFs are required. In this paper, it is suggested that 2 DOF actuators will be added to the build platform. This is a change that can easily be integrated into any conventional FDM system without much system change. In order to deposit the material, the tool head and the build platform must have cooperative control. Also, two trajectories need to be planned, one for the tool head and the other for the build platform. This paper addresses the methodology for trajectory planning to deposit any material on a curved surface with synchronization of both the tool head and the build platform. To verify the proposed method, the methodology was applied when embedding copper wire on an FDM-fabricated part resembling section of a duct or pipeline.Trajectory PlanningThis paper suggests a method for harmonic motion trajectory planning for use with an FDM extrusion head and moving build platform.

The procedure, based on robot kinematics, is as follows:1) Calculating the tangent vector of the curved surface in an object’s local coordinate system 2) Finding forward kinematics from the local coordinate 3) Finding forward kinematics from the tool coordinate to the build platform 4) Calculating joint parameters using inverse kinematics and the step 2 transformation to make the surface parallel with the FDM tool head 5) Calculating inverse kinematics and the step 3 transformation to follow the change in target position from Step 4 Deriving the tangent vector on the curved surface The first step in the process, is finding the required rotation angle to allow the tool head approach a build surface in a perpendicular fashion.

Any point on the surface of a solid object in ordinary 3D Euclidean space R3, can be generalized by a parametric surface.P3p ii 1  m u v ( , )p i i P S u v ( , )(1)where, basis i on the curve are described by parametric surface S .is the position vector for any point on the surface.P and its norm is mapped by the bounded surface parameterip is the codirectional vector with . Therefore, all positions,u vAs a 3D model’s surface is already known, based on the CAD created in Cartesian coordinates, this paper uses x, y and z as the basis of a coordinate system. Also, it is able to define surface S as the height of position x, y.YtXtZtltxc,yZ LY LX LZsx,ycY mZ mX mlmYsXsFig1. Parametric surface and coordinateszS x y ( ,)(2)Fig (1) shows an arbitrary curved surface in space. Subscript indicate the fields of the movable stage, wire embedding tool, object being built and the local coordinates, and respectively.

Subscript indicate the fields of the movable stage, wire embedding tool, object being built and the local coordinates, and respectively. On the surface, two isoperimetric curves were made parallel with each axis: axis . An origin of local coordinates was set at the intersection point of the curves and the mY was set along the tangent of the ym constant isoperimetric curve. The was set in the same constant. Finally, the axes of the local coordinate frame can be obtained by fashion at the partial derivative.,,,StmLLYmxmX LXXLS x y ( ,  x ),Y LS x y ( ,  y ),ZLn  X XLL Y L Y L(3)By neglecting the higher order terms, the first order approximation can be used to obtain the.

slope angles of tangent vector aboutandLXLYarctanS x ( , ) x y x S x y ( ,) ( , if S x y)x i2iS x y ( ,)(x )i(4)arctanS x y ( ,y )    y S x y ( ,) ( , if S x y)y 2iS x y ( ,)(y )i(5)After all, all points on the curved surface can be expressed by position and surface vectors, which can be shown as slope angles.Kinematic Modeling The end-effector coordinates of a Multi-DOF system, which consists of a number of actuators, varies depending on the movement of each actuator. Therefore, one must find the kinematics from the base to the end-effector, based on whole displacement of the actuator in the joints. This paper uses the Denavit-Hartenberg (D-H) convention to solve the kinematic problem in the spatial kinematic chain, such as in the articulated robot [7].

This paper uses the Denavit-Hartenberg (D-H) convention to solve the kinematic problem in the spatial kinematic chain, such as in the articulated robot [7]. In the D-H representation, each link in the serial chain, shown in Fig (2), is described by a coordinate transformation from the previous coordinate system, forming a homogeneous transformation, D-H matrix, which is a product of four basic transformations.A iRot Trans Trans Rot , z diz, i x ai ,x, i       i sin  i 0 0cossincos   ii cos   i i sin  i 0cossinsin   i sin   i i cos  i 0icosa i a icossin d i 1 i  i     (6)The final forward kinematics can be obtained by multiplying each transformation matrix consecutively. Conventional AM systems typically consist of just three prismatic joints to move with three DOFs in a Cartesian coordinate system.

Conventional AM systems typically consist of just three prismatic joints to move with three DOFs in a Cartesian coordinate system. Therefore, a submatrix, top left, was simplified to include constant terms in each joint coordinate transformation. This paper adopted revolute joints in the build platform to rotate an object as it is being built. The rotational transformation is shown in the kinematics between the object and build platform.3 3Trajectory Planning According to forward kinematics, the position and orientation of the end-effector in the initial frame are given by0 H T  nA q 1( 1)A q ( n n)(7)is the actuator variable of the where, can represent the desired position and orientation of the end-effector, i which will embed the wire on a surface. Next, one must find the values for the joint variables . This inverse kinematics problem is complex and becomes q 1, more difficult with increasing number of joints and complexity of joint composition.

This inverse kinematics problem is complex and becomes q 1, more difficult with increasing number of joints and complexity of joint composition. However, it is simple to solve the reverse kinematics problem because the rotational term in the transformation becomes a symmetric matrix. The joint variable of AM system will automatically follow the position of the target. If there is no redundancy in DOF and each joint is orthogonalis D-H matrix andis homogeneous transformation,iAiqHth joint. Therefore,Hq so that,0 T q 1( n,,q)Hnnon the moving platform, the variable will represent the normal vector of target also. If some redundancy in DOFs and a complex geometry exists, it can solved by jacobian pseudo-inverse[8].Fig2.

If some redundancy in DOFs and a complex geometry exists, it can solved by jacobian pseudo-inverse[8].Fig2. D-H frame assignmentApplicationFor verification of the proposed trajectory planning method, this chapter utilized the method for embedding copper wire along a curved surface.Surface Analysis Fig (3) shows a cross-sectional diagram of the duct in which a copper wire was embedded. It is of interest to embed copper wires on ducts, pipelines, and other parts to manage the electrostatic charge that can accumulate on nonconductive thermoplastic parts. For example, through the process of selective chemical adsorption, electrostatic charge can accumulate on fuel-carrying plastic pipelines that present ignition hazards due to electrostatic brush discharges and spark discharges when the pipes interact with connecting manifolds or valve bodies [9].

Static electricity can also accumulate on nonconductive plastics used as transfer containers in powder processes which can discharge and cause explosions [10]. As such, a path for dissipating electrostatic charges to earth is required in these types of parts. The wire embedding process demonstrated in this work can ultimately produce a network of conductive paths to mitigate charge accumulation that may lead to an incendive discharge.The surrounding surface of the duct is a composite shape with curves and lines. Because the cross-sectional shape is constant along the center axis, like a rod, the normal vector of the surface is just dependent of the polar angle of the center. Fig (4) shows the angle of tangential vector according to the polar angle. To embed wire around the duct, the tangent angle of each polar position must be horizontal, or perpendicular to the surface, by rotating on a center axis.Fig3.

To embed wire around the duct, the tangent angle of each polar position must be horizontal, or perpendicular to the surface, by rotating on a center axis.Fig3. cross-sectional diagram of air flow ductAngle of Surface tangent vector050) e e r g e D(100150t n e g n a T200f oe g n Al250300350050100150200250300350Polar angle (degree)Fig4. tangent vector angle of surface according to polar angleFig5. Wire embedding system and coordinate systemsSystem Kinematicsotdenotes the Fig (5) shows the wire embedding system and several coordinate systems. coordinate of the tool base and and represent the tool end-effector and movable state, respectively. The ultrasonic horn attached to the tool head of the CNC router is able to move in accordance with a 3DOF Cartesian coordinate system, just like a conventional FDM system.

The ultrasonic horn attached to the tool head of the CNC router is able to move in accordance with a 3DOF Cartesian coordinate system, just like a conventional FDM system. When the ultrasonic horn is running, copper wire is supplied through a machined hole, on the side of the horn, and fed through the tip where the copper wire is embedded through ultrasonic welding. A rotary stage was installed on the base platform of the CNC router and designed to rotate the workpiece during a build.tsTable (1) shows DH-parameters of a 3 DOF CNC router based on the coordinate frame shown in Fig (6). Because of Cartesian characteristics, each joint variable of the CNC router is separated, non-interactive and simply derived as Eq (8).Table 1. DH- parameter of CNC routerlinkdaαθ12i**d1d200*0*0π/2π/2π/2ππ/2π/2T     1 000 1 0 i0 0 1 i 00 0d2 d 1d3 1     (8)3d30*0π/2.

it can be set as 0 because of zero setting function of CNC router. **. i denotes an intermedium link.Fig 6. coordinate frame of CNC routerInverse Kinematics Two inverse kinematics must be solved so that two moving systems can be driven together harmoniously. Each point of the designated tool path, which draws a shape on the surface,becomes a target point to solve the inverse kinematics. It is recommended that the system with more movement constraints be selected as the base and its inverse kinematics be solved first. In the practice system, the rotary table just had a rotational DOF and was unable to complete translations. If an object has a non-uniform radius shape, the rotation of the object will cause the trajectory point to shift. Therefore, spatial movement,occurring from rotating the workpiece and varying its radius, should be covered by 3 DOF tool.

Therefore, spatial movement,occurring from rotating the workpiece and varying its radius, should be covered by 3 DOF tool. As has been mentioned in the surface analysis section, the rotary stage has to rotate the part to maintain the tangent of the surface horizontal. Therefore, the required angle, solution of inverse kinematic of rotary stage, is opposite-signed to that of the angle of the surface tangent vector.The motion system used in conventional AM is a 3 DOF gantry type Cartesian movement, which is called PPP (prismatic- prismatic- prismatic) type robot and its joint value is uniquely determineddirectly from inverse kinematics.Trajectory planning After solving inverse kinematics, each point of embedding path results in several trajectories, which are thedisplacement and angular position of each joint. Fig (7) shows the resulting trajectories of each joint according to time. The top graphdisplays angular trajectory of the rotary stage on which the part being built is held.

Fig (7) shows the resulting trajectories of each joint according to time. The top graphdisplays angular trajectory of the rotary stage on which the part being built is held. This is flipped from the shape shown in Fig(4).The middle and bottom graphs show trajectories of X axis and Z axis joints of tools, respectively. X axis movement is simply a repeatedforward and backward movement, but Z axis movement indicates high changes in specific points.) e e r g e d (300200n o i t a t o R100001020304050 Time (sec)6070809010020)m m10t n o Ji0s x a - Xi102001020304050 Time (sec)6070809010034)m m32(t n o Ji30s x a - Zi28262401020304050 Time (sec)60708090100Fig7. Trajectories of X,Z axis joint and rotary tableFig (8) shows the simulation results, which plotted consecutive figures at 8-second intervals. The red circle is the tool’s end position (indicated by an arrow).

The red circle is the tool’s end position (indicated by an arrow). Simulation test results showed that the tool remained on the surface continuously and the tool and part were moving harmoniously.Fig8. Simulation test (time step: 8 sec)Experimental TestThe air flow duct designed in Fig (3) was built by a Stratasys FORTUS 400mc (Eden Prairie, MN) system with polycarbonate. The tool motion of the ultrasonic horn was controlled by a Techno CNC (Ronkonkoma, NY) controller and programmed by SAC script,which is a Techno controller inherent functional library. A rotary stage, movable build platform, was operated bya ATMEGA 128 micro controller. Motion synchronization between the two systems was achieved by handshake signaling trough data ports.ig (9) shows consecutive movements of embedding copper wire around the air flow duct surface and Fig (10) shows a finished result.

Althougha small slip occurred at the beginningbecause of insufficient heat by tool, wire was embedded uniformly at most of the points as the tool maintained contact constantly.Fig9. consecutive operating sequencesFig10. wire embedded air flow ductConclusionThis paper displayed the method for embedding metallic material on thermoplastic structures built by FDM, to achieve multi-material hybrid AM technologies for 3D printed electronics. To dispose metals on a thermoplastic model without slicing a material, it was suggested to embed a wire on the freeform surface by moving the tool head and part simultaneously. Therefore, harmonic and uniform trajectories of each motion systems were required. This paper proposed trajectory planning based on robotics. The proposed methodology was verified by the application of embedding a copper wire on the curved surface of an airflow duct.

This paper proposed trajectory planning based on robotics. The proposed methodology was verified by the application of embedding a copper wire on the curved surface of an airflow duct. The result showed that the wire was embedded uniformly at most of the points as the toolheadmaintained constant contact.AcknowledgmentThe research presented here was conducted at The University of Texas at El Paso within the W.M. Keck Center for 3D Innovation (Keck Center). Through funding from the State of Texas Emerging Technology Fund, the Keck Center recently expanded to over 13,000 sq. ft., housing state-of-the-art facilities and equipment for additive manufacturing processes, materials, and applications. Specifically, this research was supported by funding from the National Science Foundation Louis Stokes Alliance for Minority Participation program under grant HRD-1139929 as well as the Mr. and Mrs. MacIntosh Murchison Chair I in Engineering Endowment.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCorrelating Large-Format AM Print Parameters to Fiber Length and Mechanical Performance of Reinforced Polymer CompositesAndrew Rhodes1, Roo Walker1,2, Tyler Smith1,2, John Lindahl2, Vlastimil Kunc2, Chad Duty1,21 Mechanical, Aerospace, and Biomedical Engineering, University of Tennessee, Knoxville, TN 2 Manufacturing Demonstration Facility, Oak Ridge National Laboratory, Oak Ridge, TNAbstractThis paper aims to correlate processing conditions to fiber length and mechanicalproperties in fiber-reinforced composites on the Big Area Additive Manufacturing (BAAM) system at Oak Ridge National Laboratory. The processing of fiber-reinforced composites has a significant influence on their microstructure, which dictates the properties of the final product.

The processing of fiber-reinforced composites has a significant influence on their microstructure, which dictates the properties of the final product. The effect of processing is poorly documented in systems like the BAAM, leaving significant opportunities to improve the mechanical performance of printed structures. In this work, fiber length distributions from pelletized feedstock were compared against those of specimens extruded under different processing speeds. The mechanical strength of each specimen was evaluated to correlate processing speed to fiber length and mechanical properties. Experimental results showed that fiber length decreases slightly with increasing screw speed. Mechanical performance was not found to be affected by the decrease in fiber length.

Experimental results showed that fiber length decreases slightly with increasing screw speed. Mechanical performance was not found to be affected by the decrease in fiber length. This research will guide future modifications to hardware design and print parameters to maintain fiber length and maximize mechanical performance.IntroductionReinforcing fiber is essential to large-format polymer additive manufacturing [1].However, the effect of processing conditions on fiber-reinforced composites in large-format additive manufacturing has not been well documented. The BAAM uses a single screw plasticating extruder to rapidly deposit large quantities of material in the printing process. Figure 1 shows an example diagram of a plasticating extruder [2]. Pellets are fed into the flights of a screw through a hopper, and then are melted through compression caused by the screw.

Figure 1 shows an example diagram of a plasticating extruder [2]. Pellets are fed into the flights of a screw through a hopper, and then are melted through compression caused by the screw. This creates a torturous environment for the fibers, which partially defines the mechanical properties of the extruded material. Fortunately, this process is almost identical to injection molding, where there have been numerous investigations concerning the influence of processing conditions of fiber reinforced plastics, and the resulting mechanical properties of the material. Wolf et al demonstrated that conditions which increase shear decrease average fiber length [3]. Wolf took fiber length distributions from different regions within the extruder, and under different processing conditions including screw speed, temperature, and die diameter.

Wolf took fiber length distributions from different regions within the extruder, and under different processing conditions including screw speed, temperature, and die diameter. Wolf found that gentler processing conditions (slow screw speed, higher extrusion temperature, larger die diameter) result in the preservation of the fiber length. Turkovitch et al. studied the fracture of glass fiber through the length of the extruder and at different screw speeds. The findings were the same as Wolf; fibers progressively break through extrusion, and fibers breakage increases with screw speed [4]. Yilmazer et al. examined the effects of processing conditions on fiber length771and part strength in injection molded samples. They reported that fiber length and tensile properties significantly decrease as shear rate is increased through the alteration of the screw speed or feed rate [5]. Hausnerova et al.

They reported that fiber length and tensile properties significantly decrease as shear rate is increased through the alteration of the screw speed or feed rate [5]. Hausnerova et al. studied fiber degradation during multiple extrusion cycles under a range of shear rates through a capillary rheometer. Hausnerova reported that under high shear rates the length of fibers decreases, with significant damage occurring during the first extrusion cycle [6]. In summary, processing conditions have been found to affect mechanical properties through reduction or retention of the initial fiber length distribution.Figure 1: Diagram of a plasticating extruder.This work examines the influence of screw speed on the fiber length and the resultanttensile properties using a large-format polymer additive manufacturing system.

This subject has been extensively studied in an injection molding setting, but it has not been well tested in a large- format screw extrusion additive manufacturing system like the BAAM. This study was designed under the fact that the BAAM uses a modified version of an injection molding extruder, and therefore the effect of processing conditions on fiber length and mechanical properties should follow those of the studies done in injection molding.MethodsPrinted SamplesThe material used in this study was Techmer Electrafil, a pelletized 20% carbon fiberreinforced acrylonitrile butadiene styrene (CF-ABS) feedstock used in screw extrusion processes [7].The pellets were an average of 2.5mm in diameter, 3mm long, and contained fibers with a diameter of 7μm. The pellets were dried at 80oC for at least 4 hours to remove moisture before printing. The processing condition chosen for this study was screw speed. All other conditions were held constant.

The pellets were dried at 80oC for at least 4 hours to remove moisture before printing. The processing condition chosen for this study was screw speed. All other conditions were held constant. The screw speeds tested were 80, 160, and 240 RPM, as it is generally accepted that the BAAM operates within this range. Samples were printed at 250oC with a77210.16mm (0.4”) diameter nozzle, 5.08mm (0.2)” layer height, and a 17.78mm (0.7”) bead width. Each sample set was printed as a 0.6m (2’) x 0.6m (2’) x 0.3m (1’) box Figure 2.Tensile specimens (ASTM D638 Type 1, Figure 3) were milled flat to about 6.5mm, and waterjet cut from the walls of the boxes [7]. The specimens were cut such that the load direction was parallel to the deposition direction. The samples were dried at 50oC for at least 48 h and left in a desiccant chamber at 23oC for at least 5 h (ASTM D618-B) prior to tensile testing. To ensure statistical significance, 5 specimens were measured for each condition.

To ensure statistical significance, 5 specimens were measured for each condition. Specimens that were damaged at any point were discarded.1’2’2’Figure 2: One of the printed boxes.Tensile testing of all specimens was performed on an MTS Criterion Series, Model 45with a 10kN load cell. The testing rate of 1mm/min was used for a nominal strain rate of .1mm/(mm min). Extension was measured using an MTS LX 500 Laser Extensometer. Specimens that fractured outside of the gage length were omitted from the data analysis.Figure 3: ASTM D638 tensile specimen dimensions773Fiber AnalysisThe steps of fiber measurement follow those outlined by Kunc et al. [8].1. Composite coupon isolationTo characterize the fiber length distribution within the ABS matrix, coupons of the composite were taken from the material remaining in between the nested tensile samples after water jetting.

These coupons were approximately 6mm tall and wide and cut to 6mm long, such that the weight was approximately 0.3g. The raw stock fiber length does not exceed 600μm, therefore the distance of an edge from the center of the coupon had to be greater than 600μm.2. Constrained removal of matrix materialThe composite coupon is placed in a glass vial with 15ml of acetone to dissolve the coupon. The coupon vials were put into an ultrasonic bath for at least an hour so that the fibers separated from the matrix. It is apparent that the matrix has been removed when clumps of fiber are no longer visible.3. Fiber sample isolationThe fiber sample isolation step requires isolating a portion of fibers for dispersion and imaging. The process involves suspending the fibers in the solution and then drawing out a portion of the fibers that will represent the whole. The fiber solutions were stirred on a vortex mixer at the highest setting for 20 seconds each.

The fiber solutions were stirred on a vortex mixer at the highest setting for 20 seconds each. Immediately following, a 5ml pipette with a 2mm diameter tip was used to draw out solution.4. Filament dispersionDispersion of the fibers was performed by diluting one part solution from the previous step with two parts acetone. This reduces the amount of polymer matrix in the solution that leaves a film on the fibers and microscope slides and reduces the likelihood of clumped fibers. The samples were vortex mixed a second time once they were diluted, and 3 to 4 drops were pipetted onto a microscope slide in a continuous, linear motion so that most of the slide was covered. The slides were allowed to sit until the acetone evaporated from the surface, which took no more than a few moments. This method gives good fiber dispersion, avoiding fiber clumping and fiber-poor regions.

This method gives good fiber dispersion, avoiding fiber clumping and fiber-poor regions. Additionally, each microscope slide shared roughly the same number of measurable fibers (~500 fibers) between slides created by the same solution. It was found that applying the solution to the slides in droplets promoted fiber clumping and uneven fiber dispersal. Proper dispersion is essential, since poor dispersion increases fiber overlap, and overlapping fibers cannot be measured.5. Imaging and fiber length measurementImaging was performed on a Keyence digital microscope using 50x magnification. The slides were imaged using the Keyence software’s image stitch function. This took a grid of images within a region encompassing the entire slide, and then stitched them together into a single high-resolution image.

This took a grid of images within a region encompassing the entire slide, and then stitched them together into a single high-resolution image. The fiber length distribution was determined by774manually measuring the length of 1500 fibers using ImageJ software, as the distribution stabilized at around 1000 [9].Average fiber length was calculated using two methods: number average and lengthweighted average, shown in Equation 1 and Equation 2 respectively. The number average is a representation of the number of shorter fibers, calculated by the sum of the individual fibers divided by the total number of fibers measured. The length weighted average is calculated as the sum of the individual fiber lengths squared divided by the sum of the individual fiber lengths. The length weighted average gives more importance to longer fibers.

The length weighted average gives more importance to longer fibers. Both values are often included in fiber length analyses and are useful in comparing the degree of length reduction of fiber on both sides of the distribution.𝑳𝒏 =∑ 𝑵𝒊𝒍𝒊 ∑ 𝑵𝒊Equation 1: Number average fiber length𝑳𝒏 =∑ 𝑵𝒊𝒍𝒊 ∑ 𝑵𝒊𝒍𝒊𝟐Equation 2: Length weighted average fiber lengthResults and Discussion:Fiber length distributionFigure 4 shows the fiber length distribution of the pellet feedstock. The feedstock had the highest distribution of long fibers, and the lowest distribution of short fibers of the samples. Two notable features are that the distribution is already quite broad, and that the distribution isFeedstock Fiber Length Distribution8%7%6%y c n e u q e r F5%4%3%2%1%0%>25 35557595 115 135 155 175 195 215 235 255 275 295 315 335 355 375 395 415 435 455 475 495Fiber Length (μm)Figure 4: Fiber length distribution of the unprocessed feedstock775bimodal.

This broad distribution is likely due to the manufacturing method, where the fibers were chopped and mixed using a dual screw extruder. The number and length weighted averages for the feedstock were 188μm and 253μm.The first condition tested was the 80 RPM screw speed. The resulting fiber lengthdistribution and averages show the shortening of the longer fibers and can be seen in Figure 5 and Table 1. Comparing the distribution to the feedstock, the second mode of the 80 RPM widened to encompass shorter fiber lengths. This wider mode contains the longer fibers that were more susceptible to breaking in the lowest shear rate tested. The fiber length averages tell the same story; the difference between the number average is 3%, and the length weighted average is 6%. This is evident that the longer fibers were broken more than the shorter were. The 160 RPM condition shows a surprisingly larger degree of attrition; a 17% reduction from the feedstock in both averages.

The 160 RPM condition shows a surprisingly larger degree of attrition; a 17% reduction from the feedstock in both averages. The 160 RPM fiber length distribution is now unimodal with the only mode being greater than the previous two distributions. The 240 RPM fiber length barely changes from the 160 RPM, as shown by the nearly identical distributions and averages.

The 240 RPM fiber length barely changes from the 160 RPM, as shown by the nearly identical distributions and averages. Since the fibers experienced such a large degree of attrition at 160 RPM, it is likely they reached a critical breaking point due to the shear in the extruder and will not shorten much more under theNumber average FLWeighted average FLFeedstock188μm253μm80 RPM182μm239μm160 RPM156μm211μm240 RPM158μm216μmTable 1: Fiber length average representations from the feedstock and tested screw speedsCombined Fiber Length Distributions18%16%14%y c n e u q e r F12%10%8%6%Feedstock80 RPM160 RPM240 RPM4%2%0%255075 100 125 150 175 200 225 250 275 300 325 350 375 400 425 450 475Fiber Length (μm)Figure 5: The combined fiber length distribution plot for each condition776conditions the BAAM can achieve.

Although the 240 RPM averages are slightly longer than the 160 RPM, this difference is insignificant.Tensile Properties vs. Print ParametersFigure 7 and Figure 6 exhibit the relationship between screw speed on the elasticmodulus and ultimate tensile strength respectively. A trend line of the length weighted fiber average is added to the plots to compare the relationship between fiber length and tensile properties. The figures show that there is a lack of correlation between fiber length and tensileElastic Modulus of BAAM CF-ABSElastic ModulusLength Weighted Fiber Average) a P G( E8.0 7.5 7.0 6.5 6.0 5.5 5.0 4.5 4.0 3.5 3.0 2.5 2.0 1.5 1.0 0.5 0.0300280260240220200180160140)m μ ( h t g n elr e b i F12080 RPM160 RPM240 RPMFigure 7: Elastic modulus vs.

screw speed and fiber lengthUltimate Tensile Strength of BAAM CF-ABSTensile StrengthLength Weighted Fiber Average60300280) a P M( h t g n e r t S el i s n e T5040302010260240220200180160)m μ ( h t g n elr e b i F140012080 RPM160 RPM240 RPMFigure 6: Ultimate tensile strength vs. screw speed and fiber length777properties. The differences between the elastic moduli vary but are within a standard deviation of each other. The variation in the ultimate tensile strength is even less, with only a slight difference between 160 RPM and 80/240 RPM, which have nearly identical values.Conclusion and Future WorkScrew speed was not found to have a meaningful effect on the fiber length andmechanical properties of fiber reinforced large-scale printed material under the conditions of this experiment.

The fiber length distribution of the pellet feedstock was already short at around 200μm, and so it is likely that the fibers are not long enough to break to significantly shorter lengths under the operating conditions as described. It is known that fiber length should have a positive correlation to tensile properties. However, these results show that although fiber length decreased in processing, the difference between the maximum and minimum fiber lengths (~30μm) seems to be insignificant and have little effect on the tensile properties of the final product. As a follow-up to this experiment, feedstock that has been pultruded will be investigated so that the initial fiber length distribution is relatively long and uniform. This will provide an analog for the degree of fiber length attrition under large-format additive manufacturing.

This will provide an analog for the degree of fiber length attrition under large-format additive manufacturing. It is now known that materials containing short, chopped carbon fiber reinforcements, like the one used in this study, are functionally unaffected by screw speed in a system like the BAAM.AcknowledgementsResearch was sponsored by the University of Tennessee and the National Science Foundation under grant number 2055529. This work was also sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle, LLC. Thanks to Techmer PM for supplying the print material, and Cincinnati Inc. for use of the BAAM.References[1] L.J. Love, V. Kunc, O. Rios, C.E. Duty, A.M. Elliott, B.K. Post, R.J. Smith, C.A. Blue, The importance of carbon fiber to polymer additive manufacturing, J. Mater. Res. 29 (2014) 1893–1898. https://doi.org/10.1557/jmr.2014.212.[2] J.

Materials Science and Technology Division, Oak Ridge National Laboratory, Oak Ridge, TN, USA c. Energy Science and Transportation Division, Oak Ridge National Laboratory, Oak Ridge, TN, USA d. Computational Sciences & Engineering Division, Oak Ridge National Laboratory, Oak Ridge, TN, USACorresponding author: leey@ornl.gov Keywords: LMD-wire, Distortion, Inter-layer time, Finite element method, Additive manufacturing, Plate oscillation, Ti-6Al-4VAbstractLaser metal deposition with wire (LMD-w) is one of the emerging additivemanufacturing (AM) technologies for large-scale aerospace components due to high deposition rates and material efficiency. However, it often results in undesired stresses and distortions due to non-uniform expansion and contraction of material during printing. Controlling inter-layer time, preheating, and clamping are the effective methods to mitigate the thermally induced stress and deformation.

Controlling inter-layer time, preheating, and clamping are the effective methods to mitigate the thermally induced stress and deformation. In this study, the effect of inter-layer cooling time on part distortion is investigated using a finite element method (FEM). The model accounts for actual tool paths, power, and cooling conditions. The results show that the model effectively captures the fluctuation of the Ti-6Al-4V plate during printing. Also, it shows an asymmetric distortion on the plate edges. Ultimately, the sequentially coupled thermal-stress simulation provided a quantitative understanding of the inter-layer cooling time on titanium plate distortion for the large-scale LMD-w process.1. IntroductionLaser metal deposition with wire (LMD-w) is one of the promising Direct Energy Deposition (DED) technologies.

IntroductionLaser metal deposition with wire (LMD-w) is one of the promising Direct Energy Deposition (DED) technologies. It is known as an ideal process for large component fabrication with moderate complexity since the wire-feeding method offers high productivity due to high deposition rates, material efficiency, and a less costly system. However, a few production challenges have not been fully resolved in large-scale LMD-w process [1].A LMD-w system generally consists of a laser heat source equipped on a robotic arm,wire feeder, and accessorial systems (e.g., tables, shielding gas, and recording tools such as cameras and thermocouples). The laser and wire feeder are moved using an automatically controlled robotic arm. The laser heat source melts the substrate, and the metal wire feeder forms This manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy.

DE-AC05-00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes. The Department of Energy will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).606a new deposit on the substrate. During printing, the material experiences multiple heating and cooling cycles. This induces thermal stress and distortion due to non-uniform expansion and contraction of the material because of the repeated heating and cooling.

This induces thermal stress and distortion due to non-uniform expansion and contraction of the material because of the repeated heating and cooling. It also contributes to form residual stress, resulting in significant distortion after the part is unclamped.Many researches have studied to find out an optimum control strategy for thermallyinduced stress, residual stress, and resultant distortion in the large part [2, 3]. Mughal et al. [4] investigated the effect of inter-layer time on the thermal stress and deformation using 2-D thermo-mechanical model. It was found that employing inter-layer cooling can reduce the part’s deformation. Zhao [5] et al. also showed that both inter-layer stress and residual stress can be reduced by controlling inter-layer cooling time. Aggarangsi and Beuth [6] reported that uniform preheating reduces residual stress by a maximum of 18%.

Aggarangsi and Beuth [6] reported that uniform preheating reduces residual stress by a maximum of 18%. However, the inter-layer cooling time decreases the production rate, while preheating delivers excessive heat to the deposit particularly in continuous printing, resulting in undesirable remelting and poor surface finish. This indicates that a compromise should be made to find optimum process parameters.Numerical simulations have been reported for large-scale AM processes. Ding [7, 8]developed a steady state Finite Element (FE) simulation to increase computational efficiency for the metal Big Area Additive Manufacturing (mBAAM) process. They showed that stress across thedeposit is uniform, but the significant residual stress develops after unclamping, which leads to large distortion. Montevecchi et al. [9] developed an FE simulation for mBAAM accounting for actual power distribution between filler wire and substrate material.

They successfully demonstrated a feasibility of the numerical modeling for a complex large excavator arm. Hence, an emphasis on AM process control has been required to mitigate (or eliminate) part distortion in large-scale titanium components including inter-layer time.The aim of the study presented below is to demonstrate part scale heat transfer simulationin an LMD-w system and structural corresponds to large-scale part distortion. One of the key aspects of this work is to account for the actual tool path, power, and cooling conditions. The thermo-mechanical model is validated with in-situ temperature and layer by layer distortion measurement.2. Experimentation6072.1. Laser metal deposition with wireThe experiment was carried out with a six-axis KUKA robot system. A fiber laser with a maximum power of 20.0 kW was used as the heat source for the LMD-w system. The laser was delivered to the work piece through an optical fiber. The laser spot diameter was 7.4 mm.

A fiber laser with a maximum power of 20.0 kW was used as the heat source for the LMD-w system. The laser was delivered to the work piece through an optical fiber. The laser spot diameter was 7.4 mm. The deposited material was fed from 1.6 mm Ti-6Al-4V wire with an Abicor Binzel wire feeder. The experiments were conducted in a closed argon environment with less than 250 ppm oxygen content. A laser line scanner was used to measure the bead height and plate distortion.2.2. Experimental configuration and observation of the processConfiguration of LMD-w experiment consisted of a Ti wall, Ti plate, ceramic washer,steel plate, and Al plate as seen in Fig. 1(a). The wall height is11.0 mm for ten-layers of deposit. The Ti build plate was mounted on the ceramic washer with 2.54 mm thickness and fastened with the steel plate using medium carbon steel bolts with the washer. The bead width was 11.5 mm, and the double bead deposit was 9 mm center to center as seen in Fig. 1(b).

The bead width was 11.5 mm, and the double bead deposit was 9 mm center to center as seen in Fig. 1(b). The Ti plate’s dimensions were 500 mm (length) x 155 mm (width) x 6.4 mm (thickness) as seen in Fig. 1(c). The thermocouples were placed with 100 mm spacing from each other on the Ti build plate marked as red spots in Fig. 1(c). The plate scan was implemented in every layer to measure the variation of the plate distortion. The measured plate distortion along A-B in Fig. 1(c) was used to validate simulation. Unidirectional scan strategy was used to create the 200 mm long deposits. The LMD-w station is shown in Fig. 1(d).Figure 1. Build geometry: the LMD-w station consisted of Ti wall, Ti build plate, ceramic washer, steel plate, and Al plate. A bolt clamp was used to fix the Ti build plate to the steel plate.Figure 2 shows the variation of laser power as a function of time during the build.

A bolt clamp was used to fix the Ti build plate to the steel plate.Figure 2 shows the variation of laser power as a function of time during the build. Thepower surges as the laser beam is powered on and is maintained for approximately 42s to create608the first bead. Then, it returns to the start position for about 8s. While the robot head moves to the start position, the laser power surges instantaneously at after 2,330s and cuts the wire to maintain the constant wire length. The velocity of laser deaccelerates to zero as it approaches the edge where the first deposit occurs. Meanwhile, the wire keeps stretching even at zero velocity. Consequently, the wire length becomes longer than desired. Then, the beam is turned on for the second bead. The inter-layer time is estimated from the time interval between the second beam- off and third beam-on as illustrated in Fig. 2.Figure 2.

Then, the beam is turned on for the second bead. The inter-layer time is estimated from the time interval between the second beam- off and third beam-on as illustrated in Fig. 2.Figure 2. Estimation of inter-layer and bead time from the measurement of laser power.Three different experimental settings were explored to investigate the relationshipbetween inter-layer time and temperature profile. Build 1 was built with clamps attached all around the substrate with no pause time between layers. Build 2 was built with clamps attached all around the substrate, but a pause for about 30 minutes was taken after the fifth layer. Build 3 was built with only center clamps attached to the substrate and without pausing the laser. The estimated inter-layer time is plotted in Fig. 3 for the builds 1-3. The inter-layer time significantly varies with the number of layers and the build settings. The inter-layer time is set in the range of 91 to 200s.

3 for the builds 1-3. The inter-layer time significantly varies with the number of layers and the build settings. The inter-layer time is set in the range of 91 to 200s. Generally, the inter-layer time is shorter at layers 1-5 and becomes longer at layers 6- 10. This occurs because the temperature increases as the deposit builds up due to a characteristic thermal feature of thin walls. In other words, the direction of heat dissipation becomes dominant along the build direction, which causes the build temperature to increase as more layers are deposited on the plate. Therefore, longer inter-layer time is recommended for constant layer temperature.609Figure 3 Estimated inter-layer time at (a) build 1 (=all clamps & no pause) (b) build 2 (=all clamps & a pause) (c) build 3 (=center clamps & no pause).In Fig.

4, the temperature profile is measured at TC-10, 15, and 18 in builds 1-3.Temperature at the center (=TC-15) is approximately 100~130 oC higher than that at the left (=TC-18) and right (=TC-10) regions. In Fig. 4(d)-(f), there are two peak temperatures observed. These peak temperatures occur because there is about a 30-minute pause right after the 5th layer is deposited. The pause seems to significantly decrease overall temperature at the TCs. The peak temperature at the center of build 2 in Fig. 4(e) is about 80 oC lower than the peak temperatures of the same center region in Fig. 4(b) and (h). It was observed that the temperature profile is not symmetric on the left and right side of the plate. The salient, non-uniformity was found at the center clamped condition in Fig. (g) and (i). The difference in temperature is about 30 oC in build 3, whereas it is about 18 oC in the builds of 1 and 2.

The salient, non-uniformity was found at the center clamped condition in Fig. (g) and (i). The difference in temperature is about 30 oC in build 3, whereas it is about 18 oC in the builds of 1 and 2. This implies that the bolt clamp leads to improved uniformity in temperature on the plate.610Figure 4. Measured temperature profile at TC-18, TC-15 and TC-10 in Fig. 1(c), (a)-(c): build 1 (=all clamp & no pause), (d)-(f): build 2 (center clamp & a pause) and (g)-(i) build 3 (=center clamp & no pause).3. Model DevelopmentThe 3-D simulation was implemented based on the numerical solution of heat transferequations. The governing equation was discretized and solved using commercial software, Abaqus [12], for heat transfer and stress analyses. The model was sequentially coupled with heat transfer-stress analysis such that it assumed that the temperature field was not influenced by phase transformation or mechanical deformation.

The model was sequentially coupled with heat transfer-stress analysis such that it assumed that the temperature field was not influenced by phase transformation or mechanical deformation. The heat transfer simulation was first performed, and the transient temperature field obtained from the analysis was applied as a thermal load for the stress analysis. A progressive material activation method was used to describe the material addition of the wire. The Ti wall was inactive at the beginning of the simulation. Then, as the material deposition started, the elements were activated according to the actual path of the beam and wire. The laser power was integrated over time and beam travel distance in a specified time step, and then the integrated flux was incorporated into the active elements. The key and detailed features of the simulation were addressed in the earlier literature [11, 13].3.1.

The key and detailed features of the simulation were addressed in the earlier literature [11, 13].3.1. Energy Balance in Laser, Wire, Deposit, and SubstrateThe laser power and wire addition determine the amount of energy input into the process and consequently affect the bead geometry. For instance, the higher power results in the wider width of a deposited bead. Also, heating the wire affects the stability of the process. Therefore, a quantitative evaluation of the energy during the LMD-w process is necessary to determine the611amount of energy coming from each wire (=ELaser) and laser (=EWire). The energy balance equations are given below.Figure 5. Schematic of laser metal deposition with wire (LMD-w) process.𝐸𝐸𝐿𝐿𝐿𝐿𝐿𝐿𝐿𝐿𝐿𝐿 + 𝐸𝐸𝑊𝑊𝑊𝑊𝐿𝐿𝐿𝐿 = 𝐸𝐸𝐷𝐷𝐿𝐿𝐷𝐷𝐷𝐷𝐿𝐿𝑊𝑊𝐷𝐷 + 𝐸𝐸𝑆𝑆𝑆𝑆𝑆𝑆𝐿𝐿𝐷𝐷𝐿𝐿𝐿𝐿𝐷𝐷𝐿𝐿 + 𝐸𝐸𝐿𝐿𝐷𝐷𝐿𝐿𝐿𝐿 𝐸𝐸𝐿𝐿𝐷𝐷𝐿𝐿𝐿𝐿 = 𝐸𝐸𝑅𝑅𝐿𝐿𝑅𝑅. + 𝐸𝐸𝐶𝐶𝐷𝐷𝐶𝐶𝐶𝐶. The input energy stems from laser and resistance heating at the wire.

+ 𝐸𝐸𝐶𝐶𝐷𝐷𝐶𝐶𝐶𝐶. The input energy stems from laser and resistance heating at the wire. The generated energy is consumed to melt the substrate (=ESubstrate) and create a deposit (=EDeposit) on the substrate, as illustrated in Fig. 5. Some energy is lost (=ELoss) by convection (=EConv.) and radiation (=ERad.) through the air given in equation 2. Based on energy conservation, the energy generated from the wire can be calculated by equation 3. A temperature rise of wire is calculated by equation 4 [14]. The predicted value can be used as a simulation input parameter. In actual printing, the wire feed rate is actively controlled to maintain a constant deposit height. For computational simplicity, the heat generated by resistance heating is incorporated into the laser heat input.

For computational simplicity, the heat generated by resistance heating is incorporated into the laser heat input. The beam efficiency value is approximated from the values in the literature [11, 15, 16], and the efficiency of 70% is selected for the simulation.𝑇𝑇1𝐸𝐸𝑤𝑤𝑊𝑊𝐿𝐿𝐿𝐿 = 𝐼𝐼2𝑅𝑅 ∙ 𝑡𝑡 = �� 𝐶𝐶𝐷𝐷 / 𝑇𝑇0𝑑𝑑𝑑𝑑� ∙ 𝑚𝑚22∆𝑑𝑑𝑤𝑤𝑊𝑊𝐿𝐿𝐿𝐿 = (𝜌𝜌𝐿𝐿(𝑣𝑣𝑤𝑤𝑊𝑊𝐿𝐿𝐿𝐿 ∙ (𝐴𝐴𝑤𝑤𝑊𝑊𝐿𝐿𝐿𝐿)𝐶𝐶𝐷𝐷 ∙ 𝜌𝜌𝑅𝑅𝐿𝐿𝐶𝐶)𝐼𝐼𝑙𝑙)where Ewire is the energy generated by the resistance heating of wire, I is current, R is the resistance of the wire, t is time, Cp is specific heat, T is temperature, is the wire heating extension,,, andis the feed rate, area, and density of the wire, respectively.𝑙𝑙𝑣𝑣𝑤𝑤𝑊𝑊𝐿𝐿𝐿𝐿𝐴𝐴𝑤𝑤𝑊𝑊𝐿𝐿𝐿𝐿𝜌𝜌𝑅𝑅𝐿𝐿𝐶𝐶During the LMD-w process, the laser beam is used as a thermal energy heat source tofuse the wire and substrate. The laser beam can be considered as a point-concentrated or volumetric ellipsoidal heat source model.

The laser beam can be considered as a point-concentrated or volumetric ellipsoidal heat source model. In this study, to accelerate the computation speed, the point-concentrated heat source is used as a heat input. It is reported that the influence of the heat612(1)(2)(3)(4)source type was not strong on the part scale temperature distribution [11]. The mathematic equation of the laser heat source is expressed by a given equation 5.(5)222(𝑥𝑥 + 𝑣𝑣𝑥𝑥𝑡𝑡) 𝑎𝑎2𝑃𝑃𝑃𝑃𝑦𝑦 𝑎𝑎𝑧𝑧 𝑎𝑎𝑄𝑄̇ (𝑥𝑥, 𝑦𝑦, 𝑧𝑧, 𝑡𝑡) =exp (−)+ is the beam absorptance, and a, b, and c are the dimensions of heat+𝑎𝑎𝑎𝑎𝑎𝑎𝑎𝑎√𝑎𝑎where P is the laser power, source along x, y, and z axis, respectively.is the velocity of moving the heat source.𝑃𝑃3.2. Boundary Conditions and Build Geometry𝑣𝑣𝑥𝑥The build geometry shown in Fig. 1 was discretized and meshed for the FEM simulation.The computation domain was created using mesh and a geometry generation toolkit CUBIT by Sandia National Laboratory.

1 was discretized and meshed for the FEM simulation.The computation domain was created using mesh and a geometry generation toolkit CUBIT by Sandia National Laboratory. Fig. 6 shows the geometry used in the simulation. A different mesh size of 1.1 mm for the Ti plate and wall, and 5.6 mm for steel plate were used in the model for computational efficiency. Hexahedral elements with DC3D8 and C3D8R types were used for thermal and structural simulation. Since two ceramic washers were used to clamp the plates, the top and bottom surface of the washer were fully constrained in the x-, y-, and z-directions during the simulation.Although a ceramic washer was used to minimize heat loss by conduction through thesteel plate, heat conduction still occurred through the contact of the washer and the gap between the Ti-plate and steel plate. The close thermal contact is accounted for as a gap conductance condition in the simulation.

The close thermal contact is accounted for as a gap conductance condition in the simulation. The extent of this region relies on the build constraint conditions, temperature, and evolving deformation of the plate [11, 17, 18]. The values of 0.4~0.8 mW/mm2·oC at the area around the washers and 0.02~0.09 mW/mm2·oC at the other regions were used as calibration parameters in this study.A combined heat transfer coefficient of convection and radiation was used for heat lossthrough the surrounding environment [19]. This equation is expressed below(6)where−4 is the combined heat transfer coefficient andℎ𝑐𝑐𝐷𝐷𝑐𝑐𝑆𝑆 = 𝜀𝜀 ∙ 24.1 × 101.61𝑑𝑑 is emissivity.ℎ𝑐𝑐𝐷𝐷𝑐𝑐𝑆𝑆 A simple isotropic plasticity model was used for the mechanical response. The thermo- mechanical properties were obtained from a material simulation tool called JMatPro [20] and additional literature [21, 22]. Homogeneous and temperature-dependent material properties used in the simulation are given in the appendix.

Homogeneous and temperature-dependent material properties used in the simulation are given in the appendix. The values are listed in Table 1-3 for Ti-6Al-4V, SS416, and a ceramic washer, respectively.𝜀𝜀613Figure 6. Build geometry of center clamped part, build 3.4. ResultsAn accurate prediction of temperature profile results in better prediction in distortion since the structural analysis is sequentially coupled with the heat transfer analysis. Figure 7 shows a comparison of the predicted temperature profiles with the measured ones. The temperature values were extracted from the model time histories at the monitoring locations of T10, 15, and 18 as shown in Fig. 1(c). Generally, a higher temperature is observed at the center region (=TC15) than the temperature at the left region (=TC18) and right region (=TC10). The peak temperatures at the center of Fig. 7(b) continuously increase as the build is deposited, while the peaks at the left in Fig. 7(a) and right in Fig.

The peak temperatures at the center of Fig. 7(b) continuously increase as the build is deposited, while the peaks at the left in Fig. 7(a) and right in Fig. 7(c) are saturated at around 1,700s~2,000s of build time. It seems that the rate of heat dissipation by conduction through the center and the sides is not identical due to the different contact conditions between the Ti plate, ceramic washer, and steel plate. Notice that the Ti plate is tied with the steel plate using steel bolts. The heat conduction rate is significantly higher at the bolts. Also, the bolts give an additional pressure to the Ti plate, washer, and steel plate. It reduces an air gap between the components and consequently leads to an increase of heat transfer through the washers. More energy is extracted from the deposit to the plate through the center, but the plate conductivity is the same. Therefore, more heat accumulated at the center plate resulting in higher temperatures.

Therefore, more heat accumulated at the center plate resulting in higher temperatures. The non-uniform cooling rate between the center and side regions leads to plate distortion or warping as seen in Fig. 8.614Figure 7. Validation of the predicted temperature profile with the measured data at the (a) left (=TC18), (b) center (=TC15) and (c) right (=TC10) region of the part.The total displacement (=U) is shown in Fig. 8 at the tenth layer during cooling. It clearly shows that the Ti plate undergoes overall upward bending at the edges. The distortion features in Fig. 8 have been exaggerated by three times for visual clarity.Figure 8. Total displacement of center-clamped and no-pause build at the 10th layer during cooling (Exaggerated by three times for visual clarity).615The predicted displacement is compared with the measured data at each layer in Fig 9.The plate scan was performed along the line of A-B as shown in Fig. 1(c).

1(c). The red and blue lines indicate the predicted and measured data, respectively. Both experimental measurements and predicted results show asymmetric upward bending. For instance, in Figs. 9(g)-(h), the displacement value on the left is larger than the right by approximately 10%. Notice that the displacement surges by about 20% between 6th and 7th layer. The longest inter-layer time is found at the 6th layer. There are some off values observed in the prediction. However, the displacement versus distance is generally agreed to be well with the experiment, which proves the accuracy of the proposed FEM model approach.Figure 9. Validation of the predicted total displacement with the experimental data at (a) layer 1, (b) layer 2, (c) layer 3, (d) layer 4, (e) layer 5, (f) layer 6, (g) layer 7, (h) layer 8, (i) layer 9 and (j) layer 10, the displacement is measured along the scan line of A-B in Fig. 1.Figure 10 shows the plate distortion during a heating and cooling period.

1.Figure 10 shows the plate distortion during a heating and cooling period. Initially, the Ti-plate undergoes downward bending as the material is added onto the plate. The newly created deposit is hot and expands the volume due to large thermal expansion at the high temperatures [10, 23]. The expansion on the hot deposit wall is relatively larger than the expansion on the Ti plate such that the plate bends downward. As the part cools during the interval between layers616(i.e., inter-layer time), the expanded material shrinks and causes a strong contraction in the center region. Consequently, the plate edges are lifted upward, and this results in the plate warping. The features of upward and downward bending are repeated until the end of the process. The repeated warping feature are well captured by the structural simulation in Fig. 10(a)-(f). The magnitude of total displacement is relatively smaller at layer 1, Fig.10(b) than at layer 10, Fig. 10(f). As seen in Fig.

10(a)-(f). The magnitude of total displacement is relatively smaller at layer 1, Fig.10(b) than at layer 10, Fig. 10(f). As seen in Fig. 3(c), the inter-layer time varies with the number of layers. The magnitude of distortion seems to be closely related to the inter-layer time.Figure 10. Repeated upward and downward plate bending over the build, the downward bending occurs during heating (a), (c) and (e) while the upward bending occurs during the cooling (b), (d) and (f).Figure 11 shows the variation of total displacement at the 1st, 3rd, 6th, 7th, and 10th layerduring cooling. At the center region, the displacement is minimal over the layers due to the clamping condition. The predicted maximum total displacement generally increases with increased build height.617Figure 11. Variation of max.

The predicted maximum total displacement generally increases with increased build height.617Figure 11. Variation of max. total displacement at the (a) layers 1, (b) layers 3, (c) layers 6, (d) layers 7 and (e) layers 10, the displacement generally increases over the printing process.The variation of total displacement was obtained at 100 mm (A’) and 400 mm (B’) on the line of A and B in Fig. 1(c) and compared with the predicted values for quantitative analysis. The blue and orange solid lines represent the predicted results at the monitoring locations. The red squares and green rhombuses indicate the corresponding measured values. Blue cross and black triangle denote simulation outliers, which are the maximum offset values from the experiment. Overall, the displacement increases as the more layers were deposited as shown in Fig. 12(a).

Overall, the displacement increases as the more layers were deposited as shown in Fig. 12(a). It has been agreed that the trend in the simulation of this experiment has an accuracy of 90% at the left and 87% at the right on average (the outliers are not included for the calculation in Fig. 12(b).). Notice that the displacement values are not identical on the A (=left) and B(=right) regions of the plate. This indicates that the distortion is asymmetric, and the plate bends more on the left side. It is likely attributed to the fact that the thermal conditions are not consistent due to the unidirectional scan direction.618Figure 12. Variation of total displacement at 100 mm (A’) and 400 mm (B’) on the line of A and B, upward trend in the prediction and experiment are well agreed. It shows that the distortion is not symmetric at the sides. The accuracy is approximately 90% and 87% at the left and right region.5.

It shows that the distortion is not symmetric at the sides. The accuracy is approximately 90% and 87% at the left and right region.5. Summary and ConclusionIn this paper, the sequentially coupled thermal-stress simulation was conducted for a large-scale LMD-w process to provide a quantitative understanding of the inter-layer cooling time on Ti plate distortion. Through the experimental and numerical investigation of the process, the following conclusions can be drawn:The simulation effectively captures thermally induced part oscillation of the Ti-plate during the deposition of material.The large-scale simulation shows about 90% accuracy in displacement prediction.

It indicates that the proposed simulation provides reliable results for the LMD-w process.The distortion is asymmetric at the Ti- plate edges due to different thermal conditions stemming from the unidirectional scan direction.The inter-layer time of a couple of minutes seems not to have a considerable influence on the distortion in the large-scale part.Although only a simple double track simulation is conducted, the modeling approachshows a potential to aid in understanding the distortion behavior during the large-scale process. Ongoing work includes an extension of the simple two beads to complex geometry, residual stress, and microstructure prediction. The heat transfer simulation needs to reduce the time step to precisely capture the temperature profiles at around the melting temperature of the material. Then, the model may be used to control microstructure and stress evolution to guarantee the desired mechanical properties in the LMD-w process.6.

Then, the model may be used to control microstructure and stress evolution to guarantee the desired mechanical properties in the LMD-w process.6. AcknowledgmentsThe research was sponsored by the US Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office under contract DE-AC05-61900OR22725 with UT-Battelle, LLC. This work is also supported in part by Cooperative Research and Development Agreement with GKN Aerospace, under contract No. NFE-15- 05725.7. References [1] D. Ding, Z. Pan, D. Cuiuri, H. Li, Wire-feed additive manufacturing of metal components: technologies, developments and future interests, The International Journal of Advanced Manufacturing Technology, 2015, 81, pp.465-481. [2] B.A. Szost, S. Terzi, F. Martina, D. Boisselier, A. Prytuliak, T. Pirling, M. Hofmann, D.J.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCORRELATIVE BEAM PATH AND PORE DEFECT SPACE ANALYSIS FOR MODULATED POWDER BED LASER FUSION PROCESSDeniz Sera Ertay*, Henry Ma*, Mihaela Vlasea*Multi-Scale Additive Manufacturing Laboratory, Department of Mechanical and Mechatronics Engineering, University of Waterloo, 200 University Ave. Waterloo, ON, N2L 3G1, CanadaAbstractThere are ongoing challenges in achieving full density metal parts via Laser Powder Bed fusion (LPBF). Numerous of studies have shown that the part density depends on the process parameters, the powder characteristics and the process environment conditions. The scan strategy and the interactions of scan paths at discontinuities such as borders, create regions with high probability of pore occurrence.

The scan strategy and the interactions of scan paths at discontinuities such as borders, create regions with high probability of pore occurrence. In this work, the complex relation between the defects and the toolpath at border discontinuities is investigated for a print recipe which gives >99.95% solid fraction in the core of the part. Samples are scanned by X-ray Computed Tomography (CT). The pore space was analyzed to extract the pore frequency, size, shape and location with respect to the scan path, border and contour strategies.IntroductionAdditive Manufacturing (AM) technologies offer new avenues for design of complex internal topologies and lightweight structures, and overall freedom in design with enhanced part functionality without the need for specialized tooling [1]. AM can also lower costs by reducing the design-to-fabrication cycle and consolidation of multi-part assemblies [1,2].

AM can also lower costs by reducing the design-to-fabrication cycle and consolidation of multi-part assemblies [1,2]. There is a breakeven point where the need for high-value, low-volume parts or technology-disruptive designs favor the adoption of AM approaches [3].Laser Powder Bed fusion (LPBF) is one of the main metal additive manufacturing processes, where a layer of metal powder is spread on the build plate and fused by a laser beam following a scan path through the cross-section area of the layer. In LPBF, the complex interaction between the powder and the energy input often results in porosity, microstructure inhomogeneity, and poor mechanical characteristics [4,5]. There is a detrimental impact of porosity on mechanical performance [5]. Porosity highly depends on the energy input, which is determined by the process parameters such as beam power, beam focus, scan speed, layer thickness etc. Morrow et al.

Porosity highly depends on the energy input, which is determined by the process parameters such as beam power, beam focus, scan speed, layer thickness etc. Morrow et al. studied the effect of porosity on the microstructure of the part and investigated how it affects the overall part quality [4]. Researchers have developed models, where the defect population and porosity were used as a tool to predict the mechanical properties of the part such as fatigue strength and estimated the part life [5]. As there is a direct impact of density on mechanical properties, it is important to reduce the defect population within the part and make it as dense as possible [6].The studies in literature have shown that the porosity of the parts highly depends on the energy input to the metal powder. The energy input is determined by several process parameters such as laser power, scan speed and layer thickness. These parameters are dominant in determining the energy input per time in a unite volume.

These parameters are dominant in determining the energy input per time in a unite volume. Early studies worked on understanding the relationship between the process parameters and the porosity of the part [7]. It has been shown that a272homogeneous energy input and temperature distribution can improve the density of the part [8]. Process parameters are then optimized to improve the density [9–16]. The pore defects are analyzed statistically depending on their size, frequency, morphology and distribution [17,18]. The different types of defects are grouped in three groups in terms of their generation causes: depression collapse, lateral pores, open and trapped pores [18]. Vast number of studies agree low energy input results in defects as there is not enough energy to melt the metal powder [9,11,14,16]. In a later study Cunningham et al. stated that excessive energy input also causes porosity of the part [12].

In a later study Cunningham et al. stated that excessive energy input also causes porosity of the part [12]. This is called keyhole mechanism, where materials can be ejected from the melt pool by applying an excessive energy input. The scan path defines how the energy input is distributed on the metal powder substrate. It plays an important role on the energy density and can also determine if there is an excessive or insufficient energy input to the metal powder [8]. Moreover, different scan strategies have been investigated to see the effect on the density of the part. The studies can be extended by deploying different scan strategies at the core and the skin of the part, and the interaction between these two scan strategies.In this study, the main goal is to identify the relation between defects and the scan path at the skin of the part, where discontinuities of the exposure points happen in LPBF process.

Samples are manufactured by using different scan strategies at the periphery (or border) and using a process recipe, which gives 99.95% solid fraction for Invar-alloy at the core of the part. Samples are scanned by a X-ray Computed Tomography (CT), which is a well-suited equipment for defect space analysis [13]. The defects are analyzed statistically to investigate the pore frequency, size, morphology and the location with respect to the toolpath of the laser. In this study, the group was able to identify skin (or periphery) strategies that improve the overall density by minimizing sub- surface pores. This would enhance the tensile and fatigue properties, as sub-surface defects have a high detrimental impact on part mechanical performance.Material and MethodsCylindrical samples are manufactured using Renishaw AM400, which is a LPBF system with a modulated laser. Invar-alloy metal powder is used to manufacture the samples.

Invar-alloy metal powder is used to manufacture the samples. It is a nickel- iron based alloy known for low coefficient of expansion. It has been used in temperature sensitive applications. The chemical content of the Invar-alloy powder can be seen in Table 1. The particle size distribution of the powder is found using a Retsch Camsizer X2 Optical Particle Size Analyzer with the X-Jet dispersion module. Minimum chord length of the powder particles is measured and averaged. The powder size distribution for Invar-alloy powder can be seen in Table 2.Table 1. Chemical content of Invar-alloy powder (Wt. %)ElementsFeNiMnCCoMoCrPWt. %Balance36.120.440.010.010.050.120.002Table 2. Powder size distribution of Invar-alloyD10D50D90273S0.003Average size of min. x c (μm)23.233.645.0The samples are 5 mm in diameter and 8 mm in height, specifically designed for ease in density and pore space measurements via CT scanning.

x c (μm)23.233.645.0The samples are 5 mm in diameter and 8 mm in height, specifically designed for ease in density and pore space measurements via CT scanning. The process parameters are optimized to achieve high density at the core of the part. In this study, the core parameters used in manufacturing the samples were kept constant, to isolate the occurrence of defects due to the contribution of parameters sets at the periphery of the parts. The core process parameters used for manufacturing the samples used in this study are summarized in Table 3, with a meander scan strategy, with a 67o meander rotation angle increment between subsequent layers. With these parameters, higher than 99.95% density has been achieved at the core of the samples.Table 3.

With these parameters, higher than 99.95% density has been achieved at the core of the samples.Table 3. Process parameters for the core scan strategy used to manufacture the samples investigated in this studyLaser powerExposure timePoint distance Hatch distance Layer thickness275 W70 μs60 μm60 μm30 μmFour different border scan strategies are used to observe the effect on the defect population at the periphery of parts. These strategies are (i) blocked path, (ii) border path, (iii) contour path, shown in Figure 1a and (iv) hatch compensation factor shown in Figure 1b. The blocked path (i) is used to reduce the number of scans in thin areas to a single scan. The border scan path (ii) circumferences the 2D cross-section area of the layer geometry and it is the outermost scan exposure. It is recommended to have a border scan at each layer to improve the surface structure and quality [19]. The contour scan (iii) mimics the border with a certain offset from the border.

It is recommended to have a border scan at each layer to improve the surface structure and quality [19]. The contour scan (iii) mimics the border with a certain offset from the border. The contour scan is intended to reduce the porosity between the hatch and the border by re-melting discontinuities. Hatch compensation factor (iv) is a feature of the modulated beam process, where additional points can be added at the end of the hatch lines, if the distance between the last exposure point on a hatch line and the border line is bigger than the user-defined threshold. This distance can be identified as a percentage of the hatch space used in the recipe. The scan strategy and corresponding process parameters used in manufacturing of the part periphery in this study is shown in Table 4. As it is summarized in Table 3, the laser power used at the core of the parts is 275 W. However, the laser power varies at the skin laser scan.

As it is summarized in Table 3, the laser power used at the core of the parts is 275 W. However, the laser power varies at the skin laser scan. The border power sample 11 is the same as core power and the blocked path power is 110 W. The border power the samples 01, 09, 17 and 25 is 250 W and the contour power is set at 50% of the core power. The skin scan re-melts some of the regions. In order to prevent excessive energy input the skin laser power is chosen lower than the core laser power. Additionally, samples with the same periphery scan strategies are printed, but with a core meander rotation angle of 107o and the rest of the parameters the same to see the effect on sub-surface density of the part.274Figure 1. Scan strategies used in this study: a) Border, hatch, contour, blocked path, b) hatch compensationfactorTable 4.

Scan strategies used in this study: a) Border, hatch, contour, blocked path, b) hatch compensationfactorTable 4. Scan strategies of analyzed samples (H: Hatch, B: Border, C: Contour, BP: Blocked path)Hatch Compensation0%25%50%H+B+CSample #01Sample #09Scan StrategyH+BSample #17Sample #25H+B+BPSample #11ZEISS Xradia 520 Versa 3D X-ray Nano-CT is used to scan the samples. The CT data give the 3D visualization of the part with the defects. For this scan, the voxel size of 12 μm is used to detect the defects. CT analytics were deployed to extract the coordinates of the defect centroid within each CT scanned surface and in 3D space, which determines the defect location relative to the sample geometry. The location of the defect center at each layer is extracted. The size of the defect cross-section in pixels at each layer is given to have a better understanding of how the defect is seeded, grown up and shrunk.

The size of the defect cross-section in pixels at each layer is given to have a better understanding of how the defect is seeded, grown up and shrunk. The volumetric fraction is another output of the CT data. It is given as a percentage at each layer. The cross-section size and the volumetric fraction data give a quantified measure of how the defect is seeded, grown up and shrunk. A distinct number is assigned to each individual defect within a part, to keep the track of defects at each layer.For this analysis the defects at the core of the part are not included, as the amount is negligible compared to the defects occurred at the skin of the part and were optimized in another study. It has been previously observed that for parameter sets optimized in the core, the defects exist where the core scan path discontinuities intersect the skin scan path(s).

It has been previously observed that for parameter sets optimized in the core, the defects exist where the core scan path discontinuities intersect the skin scan path(s). This means that the process recipe works well for achieving high density parts when the laser exposure points are homogeneously distributed, and that the edges of the part are problematic. In this study, the group was able to identify skin (or periphery) strategies that improve the overall density by minimizing sub-surface pores.275Figure 2. Data registration of CT with the scan pathDistinctive registration markers are printed on top of the samples, shown in Figure 2, to register the CT scan data with the scan path and to help in part identification. Only the laser exposure points at the end of the hatch lines and the skin (border, contour and blocked path) exposure points are extracted from the scan path data as this study is focused on analyzing the defects occurred at the skin of the part.

The defect centers at each layer are analyzed together with respect to the near-by laser exposure points.Statistical Analysis of the Defect SpaceThe CT scan data was analyzed statistically to determine which skin scan strategy works best for achieving high density parts. In Figure 3, the periphery defect population for all the parts manufactured with different scan strategies are shown. The most defects have been obtained in sample 11, where border and blocked path scan strategies are used at the skin of the part. The reason for that could be excessive number of laser exposure points at the skin of the part. The core meander hatch lines are rotating 67o from one layer to another in all these parts to have an alternating direction of the beam raster direction between layers and avoid compounding layer-by- layer defects resulting from stacking the beam path.

On the other hand, the skin exposure points, on border, contour and blocked path, are not rotating and changing position on the x-y plane. They occur at the same position on each layer. It means that the energy is localized on specific points. The stacked laser exposure points contribute to the defects occurring at the skin. Furthermore, the blocked path lines create more discontinuities at the skin of the samples. Although they are homogeneously distributed through the circumference of the sample, they are still causing a discontinuity radially. In this study, only the defects at skin are analyzed, where the skin is 0.2 mm deep from the actual part border radially.276Figure 3.

In this study, only the defects at skin are analyzed, where the skin is 0.2 mm deep from the actual part border radially.276Figure 3. a) Defect population in different samples, where H, B, C, BP and HC represent hatch, border, contour, blocked path and hatch compensation, respectively, b) the flattened CT images of the cylindrical part (sample 11) black spots show the defects, c) the flattened CT images of the cylindrical part (sample 25) black spots show theThe histogram of defect population in different samples can be seen in Figure 3. There is an approximately 84% reduction of defect population on the skin of the part, when the blocked path scan strategy is muted. The other four parts show a similar range of defects at the skin. The best performance is observed, when only border strategy is used on the skin and 50% hatch compensation is enabled.

The other four parts show a similar range of defects at the skin. The best performance is observed, when only border strategy is used on the skin and 50% hatch compensation is enabled. An in-depth discussion is provided by analyzing the population of defects in detail in terms of pore size and depth.The depth of the defects is analyzed statistically for each part. The depth is quantified as the number of process layers over which a pore defect exists. Depth of two means that the defect propagates through two layers of the process, where the layer thickness is 30 μm for this process recipe. The Figure 4 shows the histogram of the defect population grouped according to their depth. It can be seen that each part shows the same trend in the histograms. Most of the defects propagate through 2 and 3 layers in each part. In general, there are only a few defects, which propagates through more than 5 layers or less than 2 layers. The depth can give an idea about the cause of the defect.

In general, there are only a few defects, which propagates through more than 5 layers or less than 2 layers. The depth can give an idea about the cause of the defect. In literature it has been stated that shallow defects with irregular shape are caused by poorly chosen process parameters, where the energy input is not enough and the process is under-melting [17]. The shallow defects occurring in less than one-layer thickness are usually caused by277entrapped gas. The depth of the defect can also change during manufacturing of the part. If a defect occurs due to lack of energy, this can be compensated by the laser exposure points on the next layer. So that the next layer can play a corrective role for the previous layer.Figure 4. Histogram of the defect depth at the skin, given in number of layers. It shows the depth distribution of the defects within one part.The morphology of the defects is analyzed statistically. It can be also an indicator of the defect initiator.

It shows the depth distribution of the defects within one part.The morphology of the defects is analyzed statistically. It can be also an indicator of the defect initiator. In this analysis, the biggest volumetric fraction of the defect is located within the defect layers. If there is a defect with 4 layers in depth and the biggest volume of the defect occurs in the second layer, it means that the biggest volumetric fraction is at the second layer of the defect from the bottom. The defects are grouped according to their depth. Figure 5 shows the histograms of the biggest volumetric fraction location within the defect for each depth of the defect. It is observed that the biggest volumetric fraction usually occurred in the middle layers of the defect. It means that the defect is seeded first, it grows up and then shrinks during the manufacturing of the part. For defects which span multiple layers, the largest circumference of the defect occurred in layers 2 and 3.

For defects which span multiple layers, the largest circumference of the defect occurred in layers 2 and 3. For instance, in Sample 25, defects span up to 5 layers; for example, for the population of defects spanning 4 layers Figure 5 (25-c), the largest segment in most pores occurred in the second layer.278Figure 5. For most defects which span multiple layers, the histogram showcases the frequency of the layer where the biggest pore volume exists. The layer is indicated, where the biggest volumetric fraction of the defect happens.In the case of the samples studied, one of the main causes for a defect is lack of fusion. It is observed that there are less exposure points on the skin of the samples at specific locations, depending on the orientation of the hatch lines in relationship to the skin scan strategies. In order to determine these locations analytically, the position of the laser exposure points and the defect is transferred to polar coordinates.

In order to determine these locations analytically, the position of the laser exposure points and the defect is transferred to polar coordinates. The angle between the laser exposure point normal to the surface of the sample and the hatch lines are calculated. This angle is called α shown in Figure 6. When this angle gets closer to 90o, the distance between the laser exposure points at the end of the hatch lines increases, with less laser exposure points per unit area at these specific regions. At the regions, where this angle gets closer to 0o, the population of the laser exposure points increases. Hence, the distance between the laser exposure points gets closer to the distance between the hatch lines.279Figure 6. Angular positions of the defects, laser exposure points are shown in θd and θep. The hatch angle is shown in θh with respect to the x-axis.

Angular positions of the defects, laser exposure points are shown in θd and θep. The hatch angle is shown in θh with respect to the x-axis. The angle between laser exposure points’ angular position and the hatch angle is represented by α.The defects are investigated layer-by-layer with respect to the scan path of the laser and the laser exposure points. Figure 7 shows the hatch lines in purple, the skin scan path in dark blue and the actual border of the part in pink. The circles show the defect location at each layer and the red circle is used to highlight the biggest volumetric fraction of the defect. It can be seen that the laser exposure points are getting further away from each other when the angle between the laser exposure point normal and the hatch lines is almost perpendicular. This effect can be better observed in Figure 7 on image 25-a) and 11-c).280Figure 7.

This effect can be better observed in Figure 7 on image 25-a) and 11-c).280Figure 7. Randomly selected defects with 3 layers in depth from sample 11 and 25 and the scan path on these layers are plotted together. The red and green dots show the centroid of the defect. Red color indicates the biggest volumetric fraction of the defect.An additional set of experiments was planned to see the effect of the rotation angle of the core meander hatch lines on the defect population at the border. The samples are manufactured with the same processing parameters and the hatch rotation angle is set to 107o. The defect population at the skin of the samples decreased by 20% as shown in Table 5. A random region is observed throughout the layers of the sample 11 and 25. It can be seen that the density of the laser exposure points is changing frequently. For instance, if the density of the laser exposure points is low in one layer, it will be higher on the next layer.

For instance, if the density of the laser exposure points is low in one layer, it will be higher on the next layer. This is a result of the rotation angle in this set of experiment being closer to 90o than the first set of experiment. Hence, the layer on top with more laser exposure points is a corrective action to the previous layer. It compensates the lack of energy on the previous layer.281Table 5. Comparison of defect population between two different experiment sets with different hatch rotation angles 67o and 107oDefect Population (Number of Defects)Hatch rotation angle 67o 107oSample 0169 56Sample 0991 60Sample 1764 50Sample 2554 38It is already observed that the distribution of the laser exposure points is not homogeneous at the skin of the part.

In order to see the effect of the density of the laser exposure points on the defect population at the border, the cylindrical samples are sliced, and each slice is investigated throughout the sample at each layer shown in Figure 8. For each slice at each layer, the hatch end points are counted, and that slice is flagged if there is a defect located. With this analysis, it is aimed to develop algorithms to predict defects using a probability distribution function, when a certain set of laser exposure points occur in consecutive layers using statistical and machine learning algorithms. This is part of a work in progress and will be presented as part of published work.Figure 8.

This is part of a work in progress and will be presented as part of published work.Figure 8. Investigation of laser exposure points and defects at each slice layer-by-layerConclusionIn this study, the defects at the skin of the samples are statistically analyzed according to their size, distribution, depth and frequency on samples printed with the same process parameters in the core and different scan strategies at the skin. The statistics of defects in different samples are observed to understand the complex relation between the scan path and the defect occurrence. It is seen that the size distribution of the defects has a very similar trend in every sample although they are printed with different scan strategies. This means that the size distribution and frequency do not change according to the scan strategy. It only dampens the amount of the defects occurred in one sample.

This means that the size distribution and frequency do not change according to the scan strategy. It only dampens the amount of the defects occurred in one sample. The same behavior is also detected in the morphology of the defects, where the defect grows up and shrinks down volumetrically through the longitudinal direction. For future research directions,282a statistical or machine learning tool will be designed to predict the probability of the defect occurrence for a given scan path and strategy. So that possible defects can be detected and avoided by modifying the scan path before manufacturing the parts. This tool will also help to increase the density of the parts especially at the discontinuities of the scan lines (skin of the part) manufactured by LPBF process.AcknowledgementsThis work was supported by funding from the Federal Economic Development Agency for Sothern Ontario (FedDev Ontario), in partnership with our industry partner Burloak Technologies.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCOST IMPACT OF THE RISK OF BUILD FAILURE IN LASER SINTERINGM. Baumers* and M. Holweg†Additive Manufacturing and 3D Printing Research Group, Faculty of Engineering, University of Nottingham, Nottingham, NG7 2RD, United Kingdom †Saïd Business School, University of Oxford, Oxford, OX1 1HP, United KingdomAbstractWhile the feasibility of adopting Additive Manufacturing (AM) has been demonstrated in a range of industrial sectors, the total costs associated with the operation of the technology are not fully understood. This study reports the results of a series of build experiments in a controlled environment for the analysis of the total cost of the AM technology variant Laser Sintering (LS).

This study reports the results of a series of build experiments in a controlled environment for the analysis of the total cost of the AM technology variant Laser Sintering (LS). Incorporating a structured representation of the process flow of LS, the developed cost model shows for a LS system of the type EOSINT P100 that the expected cost impact of build failure has a substantial effect, responsible for a share of up to 38% of total costs. The analysis further demonstrates that, due to the adverse effects of such ill-structured costs, the cost efficient level of build volume utilization is sub-maximal. This result is discussed in the context of the operational reality of using LS technology and the availability of economies of scale.IntroductionThe progressing diffusion of Additive Manufacturing (AM) is providing significant scope for manufacturing innovation. The technology is regarded as an opportunity to rethink design, manufacture on demand and generate customized products.

The technology is regarded as an opportunity to rethink design, manufacture on demand and generate customized products. The operating principles behind the umbrella term AM range from the thermal fusion of material particles to the exposure to UV radiation of photoreactive monomer resins (Gibson et al., 2014). AM is associated with two generic benefits (Tuck et al., 2008): firstly, AM removes many tooling-related geometrical constraints that are relevant in conventional manufacturing processes (Hague et al., 2003). Secondly, AM permits the realisation of products in small volumes, thereby enabling the manufacture of customized and differentiated designs (Berman, 2012; Eyers et al., 2008). It should be noted, however, that AM is currently limited by a set of generic factors detrimental to the manufacturing of end-use products (Ruffo and Hague, 2007).

It should be noted, however, that AM is currently limited by a set of generic factors detrimental to the manufacturing of end-use products (Ruffo and Hague, 2007). These factors include deposition speed, materials characteristics, precision, surface quality, process repeatability, and cost at medium to high volumes. As an additional limitation, Huang et al. (2015) identified missing standards for manufacturing applications, particularly relating to material properties.This paper reports key results of the project “The Economics of 3D Printing: A Total Cost Perspective” (Baumers et al., 2016) funded in the context of the 3DP-RDM feasibility study competition (3DP-RDM, 2016). It addresses the requirement for more realistic AM cost estimation by developing a novel type of model that is able to reflect the expected cost of build failure and part rejection. Shaping the process economics of AM, both aspects have been ignored in previous cost models.

Shaping the process economics of AM, both aspects have been ignored in previous cost models. To implement the model, this paper adopts an experimental approach relying on a series of build experiments performed on a polymeric EOSINT P100 Laser Sintering (LS) system from equipment manufacturer EOS GmbH in Germany (EOS GmbH, 2015). LS selectively builds up material within a bed of unprocessed powder allowing the user to fill the available build volume1910in three dimensions. This characteristic results in a highly parallel process, permitting the manufacture of multiple components (Ruffo et al., 2006). LS is used, for example, in aerospace, automotive, medical, industrial, consumer, and of course, prototyping applications (cf. Wohlers, 2014).Background and methodologyThe goal of strategic cost management is to employ information about cost to determine business strategies for competitive advantage (Shank and Govindarajan, 1993). As demonstrated by Baumers et al.

As demonstrated by Baumers et al. (2013), AM is characterized by the absence of long and complex supply chains normally associated with global mass manufacturing processes. Therefore, AM is likely to improve transparency with regards to production inputs, aiding the development of innovative business strategies.The seminal AM cost model was implemented by Alexander et al. (1998) and belongs to the class of Activity-Based Costing approaches (cf. Niazi et al., 2006). The model is based on time estimates for multiple process steps, including the time required for the core AM build process. The framework is adaptable to different operating principles found in AM and their effect on build time. Further, Alexander et al. integrated a structure for supporting activities and associated these with pre- and post-processing costs. As noted by Son (1991), however, the cost characteristics of production technology is also shaped by ill-structured aspects.

As noted by Son (1991), however, the cost characteristics of production technology is also shaped by ill-structured aspects. In the AM context, significant ill- structured costs arise from build failure, machine setup and inventory. In an effort to draw such requirements into AM cost models, quality control systems have been incorporated (Schmid and Levy, 2014; Berumen et al., 2010). This research extends such frameworks by systematically investigating the ill-structured cost effect of build failure and part rejection.To provide a body of data for this research, a series of build experiments was carried out on an EOSINT P100 LS system, which employs a 30W CO2 laser and was configured to default (factory standard) settings with a layer thickness of 100 μm and using the standard PA2200 material. The build experiments utilized a testing geometry modified for this investigation.

The build experiments utilized a testing geometry modified for this investigation. As shown in Figure 1a, the geometry features a star shape to maintain packing density at realistic levels, aiming for a packing ratio of around 10% at full capacity utilization, as shown in Figure 1b. The build experiments contained further tensile specimens for the validation of material properties (ISO 527-2:1996, not shown in Figure 1b).1911Figure 1: Testing geometry and build volume at full capacity utilizationExploiting the entire build volume (with a usable Z-height of 330 mm), corresponding to operation at full capacity, would have necessitated extensive and costly builds. Therefore, to allow a sufficient number of repeat experiments in the available time frame, the available machine capacity was artificially restricted to a shallow horizontal band of the build volume.

Therefore, to allow a sufficient number of repeat experiments in the available time frame, the available machine capacity was artificially restricted to a shallow horizontal band of the build volume. This 30mm ‘slice’ of build space was filled using a computational build volume packing approach and the data generated during the builds were extrapolated to reflect the system’s full Z-height. Overall, 14 build experiments were executed, with 10 builds at full capacity (using the available 30 mm slice) containing 5 parts each. The remaining four build experiments reflect build configurations with sub-normal levels of build volume utilization to gather data on machine operation where the available capacity is not fully employed.As suggested by Rickenbacher et al. (2013), a central aspect of cost modelling for AM is to consider processes beyond the core manufacturing step.

(2013), a central aspect of cost modelling for AM is to consider processes beyond the core manufacturing step. Matching this view, this paper uses a structured representation of the processes required for LS, as shown in Figure 2, identifying process steps performed by the technician (white boxes) and the elements of the AM process itself (grey boxes). Outright build failure events leading to premature build termination and other types of deviation from normal operation leading to part rejection are represented by two decision nodes.1912Figure 2: A process map for polymeric Laser SinteringThe durations of the process elements occurring within the AM system (warm up, deposition processes and cool down) were recorded through a digital power meter which also captured the system’s energy consumption. The durations of individual process steps in the AM work flow outside the core system were timed by the machine operator and entered into a spreadsheet.

The durations of individual process steps in the AM work flow outside the core system were timed by the machine operator and entered into a spreadsheet. The PA 2200 raw material usage was measured by weighing the powder in the build container on digital scales.Apart from the data originating from the build experiments, the cost model required a range of data from outside sources, such as suppliers and the literature. Table 1 summarizes the collected data points and their respective data sources. The majority of cost data were adapted from previous work by Ruffo et al (2006) and Baumers et al. (2015). Data on the cost of technician labor inputs were obtained from the employer (University of Nottingham).

(2015). Data on the cost of technician labor inputs were obtained from the employer (University of Nottingham). Where cost information was obtained in a currency other than pounds sterling (£), the currency was converted using the exchange rate at the publication date.1913Table 1: Summary of external data collectedCost model element Data pointProduction overhead rateAdmin overhead rateMachine purchaseDepreciation periodIndirect costAnnual operating timeEstimated maintenance and consumablesTotal machine cost rateTotal indirect cost rate (ĊIndirect)Value Unit£4.53/ h£0.31/ h£140,5008 Years5,000 H£8,516/ year£5.22/ h£10.06/ hSource Adapted from Ruffo et al. (2006) Adapted from Ruffo et al. (2006) Adapted from Baumers et al. (2015) Adapted from Ruffo et al. (2006) Adapted from Ruffo et al. (2006) Adapted from Baumers et al. (2015) Adapted from Baumers et al. (2015) Adapted from Baumers et al.

(2015) Adapted from Ruffo et al. (2006) Adapted from Ruffo et al. (2006) Adapted from Baumers et al. (2015) Adapted from Baumers et al. (2015) Adapted from Baumers et al. (2015) -Full annual labor costs£32,420/ yearLabor costWorking days net of holidayTotal hours worked per year228 Days1653 HLabor cost rate (ĊLabor)£19.61/ hRaw material price£45.05/ kgEOS GmbH (2015)Material density, as deposited0.93 g / cm³ EOS GmbH (2015)Direct cost dataEnergy priceFixed energy consumption per build Energy consumption rate£0.02/ MJ25.23 MJ1,407.50J / sAdapted from Baumers et al. (2015) Adapted from Baumers et al. (2015) -Results and discussionIn total, 63 test parts (as shown in Figure 1a) and 56 tensile specimens were built, including an additional build to replace rejected test geometries. A nominal part volume of 2521 cm³ (including tensile coupons) was deposited.

A nominal part volume of 2521 cm³ (including tensile coupons) was deposited. This required 13.09 kg of new PA2200 powder material, the amount of material was determined by the machine operator. Since two outright build failures occurred during the experiments and one additional build was necessary to replace rejected parts, 17 build experiments were performed in total. Table 2 provides a summary of the performed experimental work.1914Table 2: Summary of identified research puzzles and formulated research hypothesesExperiment no.1234567TypePlannedPlannedPlannedPlannedPlannedPlannedPlannedCapacity utilizationFullPartialPartialPartialPartialFullFullNumber of test geometries included Total Z-height (including blank layers and tensile bars) Warm up time (incl.

blank layers and tensile bars) Warm up time (incl. deposition of blank layers and tensile bars)539.1 mm206 min539.1 mm234 min539.1 mm164 min539.1 mm233 min539.1 mm204 min539.1 mm199 min539.1 mm206 min539.1 mm165 minNet normal build time152 min152 min70 min148 min149 min151 min150 min148 minCool down time960 min900 minN/A720 min960 min1080 min960 min780 minTotal build time1318 min1286 minN/A1101 min1313 min1430 min1316 min1093 minOutright build failureNoNoYesNoNoNoNoNoNumber of parts rejected20001000Material testing failureNoNoNoNoNoNoNoNoThroughout the series of builds, two separate failure modes were observed. The first kind of failure was outright build failure in which premature stoppage leads to the termination of the build, also assuming that all parts contained in the build volume are lost due to uncontrolled deformation.

During the builds, four such outright build failures were recorded, two happening during the build experiments (as listed in Table 2) and two in adjacent builds. This (very limited) dataset allowed the estimation of a mean number of depositable layers between build failures (μ = 4040.71) and the calculation of standard deviation (σ = 3267.95). Clearly, the number of failure- free successively deposited layers N can be greater than the total number of possible layers in each individual build, as limited by build volume Z-height (330 mm, shown in Figure 1b). The results19158Repeat buildFull539.1 mm190 min150 min840 min1180 minNo1 No17Additiona l buildPartial335.6 mmNot assessedNot assessed Not assessed Not assessed No0Not assessedsuggest that undisturbed operation of the LS system is likely to span multiple consecutive builds. To be able to use these estimates in the cost model, two further simplifications are made.

To be able to use these estimates in the cost model, two further simplifications are made. Firstly, the probability of build failure is treated as a function P of N, following the normal cumulative distribution function, with distribution mean μ and standard deviation σ:𝑁𝑁2−(𝑡𝑡−𝜇𝜇) 2𝜎𝜎21𝑃𝑃(𝑁𝑁|𝜇𝜇, 𝜎𝜎) =� 𝑒𝑒 −∞𝜎𝜎√2𝜋𝜋Incorporation of this failure model in the cost model enables the modification of an otherwise deterministic Activity-Based Cost model. The second simplification is to assume that the cost impact of build failure can be characterized by multiplication of the risk free cost model with the inverse of 1- P(N | μ,σ).In the schematic of the LS process flow described in Figure 2, the effect of outright build failure is limited to the first decision node. Therefore, only the process steps preceding this node, Ri, are subject by this risk of build failure.

Therefore, only the process steps preceding this node, Ri, are subject by this risk of build failure. The following process elements, S i and Ui, remain unaffected.The second failure mode pertains to the rejection of individual parts post completion of the build, due to non-permissible deformation or surface artefacts. Such failure occurs, for example, if foreign objects disturb the deposition process without leading to machine stoppage. The technician assessed the geometries visually and dimensionally for this failure mode. Throughout the build experiments, a total of four test geometries were rejected, one of which was deemed correctable, hence only three replacement parts were manufactured.This investigation assumes that the risk of part rejection p reject is constant, which can be justified on the basis that the test parts share the same geometry. On the grounds of a planned quantity of 60 test specimens, preject was estimated at 0.07.

On the grounds of a planned quantity of 60 test specimens, preject was estimated at 0.07. This risk affects a larger number of process steps so its decision node occurs later in the schematic work flow shown in Figure 2. As done with outright build failure, this risk characteristic was inserted into the cost model by forming the inverse of its complement.Following the specification of the failure models, it is clear that the probabilistic term of outright build failure P(N | μ,σ) is treated as unrelated to the product geometry. Moreover, this term is treated as unrelated to the fixed probability of part rejection p reject, implying that build configuration does not impact preject. Both aspects are additional simplifications.The proposed cost model allocates indirect costs over build time through an indirect cost rate ĊIndirect.

Both aspects are additional simplifications.The proposed cost model allocates indirect costs over build time through an indirect cost rate ĊIndirect. Additionally, the model adds the labor costs of the process steps in the work flow (partially subject to the risks of failure), ignoring the costs of other equipment outside of the AM system. As described in the process map in Figure 2, outright build failure affects a subset of the overall AM process chain (Ri).

As the main result of this investigation, Figure 3 shows the resulting unit cost function against the number of test geometries n contained in the build volume. The ‘saw tooth’ shape of the unit cost function is a result of the extrapolation procedure and therefore carries no meaning. As can be seen, unit cost differs with the level of capacity utilization, from inefficient configurations with a few units in the build volume to a situation fully using up the available build space, as illustrated in Figure 1b. Figure 3 also identifies the level of capacity utilization at each quantity as a build volume utilization ratio which is obtained by dividing the total volume of part geometry over the volume by the useful build volume cuboid.

On LS machines, build volume utilization ratios typically range from 5% to 10% in practice.1917(3)(4)Figure 3: Relationship between unit cost and capacity utilizationHence, the total cost model proposed in this investigation suggests that (subject to the risk of build failure) unit cost decreases initially but then begins to rise slightly as capacity utilization nears its maximum. The reason for this is an increase of the expected cost of build failure overpowering cost benefits of improved capacity utilization. This can be compared with a model ignoring this risk (Figure 3, dashed line). At full capacity, at a build volume utilization ratio of 9.11%, the proposed model suggests that the specific cost of operating the investigated EOSINT P100 system amounts to £0.89 per cm³.

At full capacity, at a build volume utilization ratio of 9.11%, the proposed model suggests that the specific cost of operating the investigated EOSINT P100 system amounts to £0.89 per cm³. Due to the inclusion of the risk of build failure, however, the minimum specific cost is observed at a lower capacity utilization rate of 6.62%, at £0.84 per cm³.Further insight can be gained by breaking down the total cost into its constituent elements for both levels of capacity utilization, as done in Figure 4 and Figure 5. It is shown that the cost makeup changes with the degree of capacity utilization: the cost component resulting from the risk of build failure, which is obtained by subtracting the costs without risk from the total cost, dominates the process at full capacity utilization (£553.18 or 38%). In the cost efficient configuration, the largest cost contributor is indirect process cost (£377.16) with risk-related costs accounting for £288.22 (or 29%).

In the cost efficient configuration, the largest cost contributor is indirect process cost (£377.16) with risk-related costs accounting for £288.22 (or 29%). A further noteworthy point is that labor costs make up only a minor share of the total cost, ranging from £61.20 (or 4.20%) to £49.24 (or 4.93%), respectively.1918Figure 4: Break down of total cost at full capacity utilizationFigure 5: Break down of total cost at minimum cost operationResults and discussionThe results presented in this paper go beyond existing cost models by exploring a complex relationship between the quantities of parts included within a build volume and the resulting unit cost in AM. As expected, low machine utilization levels will result in high unit cost.

As expected, low machine utilization levels will result in high unit cost. However, at high or maximal levels of capacity utilization, the expected cost impact of build failure overpowers the efficiency gains of exploiting the available capacity.This indicates that the lowest cost configuration may not coincide with full utilization of the available build space, which explains the practice of constraining the Z-height of builds to enable a more efficient process. This aspect is reported to have an impact on the general build volume packing problem encountered in AM (Araujo et al., 2015); it also lends support to cost1919models attributing costs to parts by assessing the degree to which they protrude into horizontal bands of otherwise unfilled build space (Rickenbacher et al., 2013).

Therefore, such approaches could be used to apportion the expected financial cost of build failure.In consequence, am technology operators AM face the problem of balancing capacity utilization and risk of build failure. For machine architecture, the results reached in this paper suggest that process stability should be improved such that the expected costs of build failure no longer outweigh efficacy gains originating from the exploitation of the available Z-height. This would allow the costs of fixed process increments, such as warm up and cool down activities to be amortized over a larger number of product units, putting AM users in the position to capitalize on process-level economies of scale.References− 3DP-RDM, 2016. 3DP-RDM: Defining the research agenda for 3D printing-enabled re- distributed manufacturing. University of Cambridge. EPSRC Reference: EP/M017656/1− Alexander, P., Allen, S. and Dutta, D., 1998.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCost Modeling and Depreciation for Reused Powder Feedstocks in Powder Bed Fusion Additive ManufacturingMichael Barclift1, Sanjay Joshi2, Timothy Simpson1,2,3, and Corey Dickman31Department of Mechanical and Nuclear Engineering 2Department of Industrial and Manufacturing Engineering 3Center for Innovative Materials Processing through Direct Digital Deposition (CIMP-3D) The Pennsylvania State University, University Park, PA 16802AbstractCost modeling for Powder Fusion (PBF) has traditionally treated the material feedstock as a fixed cost. Given that a built-up geometry in PBF must be in a bed filled with surrounding powder, the material feedstock is susceptible to satellites, chemical contamination, and dissimilar properties with each subsequent reuse.

In this paper, we extend an existing PBF cost model and propose a new financial depreciation model for reused metal powders. Using Sum-of-the-Years Digits depreciation, powder feedstock is valued as a function of build cycles endured by the material feedstock. A case study is presented on two example parts in Direct Metal Laser Sintering (DMLS). Results show that cost models using a fixed material cost can undervalue build jobs with a high value virgin powder by as much as 3-11% or 13-75% depending on the material and its maximum build cycles in PBF.IntroductionAdditive Manufacturing (AM) is a process of “joining materials to make objects from 3Dmodel data, usually layer upon layer” [1]. Powder Bed Fusion (PBF) is one AM process where “thermal energy selectively fuses regions of a powder bed” to produce parts [1].

Powder Bed Fusion (PBF) is one AM process where “thermal energy selectively fuses regions of a powder bed” to produce parts [1]. As shown in Figure 1, PBF consists of a powder delivery system, where a feed bed (i.e., hopper, dispenser) supplies a layer of powder onto a part bed containing the build job geometries. Thermal energy from an electron-beam or laser scans the part bed surface and fully melts a region of powder particles to form a solidified cross-section. The powder bed lowers and a coating mechanism, (i.e., blade, rake, roller) spreads additional powder from the feed bed on top of the scanned layer. Energy is applied to the newly recoated surface, and the AM process repeats for each layer until all geometries have been fabricated. At the completion of the build job, parts are removed from the machine, while surrounding un-melted powder is recovered from the part bed and overflow bin for reuse in later builds.

At the completion of the build job, parts are removed from the machine, while surrounding un-melted powder is recovered from the part bed and overflow bin for reuse in later builds. PBF is commonly executed in a build chamber filled with inert gas (e.g., Argon, Nitrogen) or under vacuum to mitigate the reactivity of the powder feedstock during the melting of each cross-section.Technologies for PBF include Direct Metal Laser Sintering (DMLS), Selective LaserMelting (SLM), Direct Metal Laser Melting (DMLM), Direct Metal Printing (DMP), Laser Melting (LM), LaserCUSING, and Electron Beam Melting (EBM) [3]. In comparison to polymeric AM processes, PBF with metal powder feedstocks requires extensive labor in order to generate a fully-functional component. In laser-powered PBF, digital modeling of the part orientation and support structures is a non-trivial task, with the goal of anchoring [4] the part to the build substrate.

In laser-powered PBF, digital modeling of the part orientation and support structures is a non-trivial task, with the goal of anchoring [4] the part to the build substrate. This design process is important due to the risk of thermal distortion and delamination during the build [5]. In electron-beam PBF, the heated powder bed can mitigate2007thermal distortion; however, the design of the support structures may need to be conducted in a manner to dissipate heat or minimize curling for fine features [6]. Finally, after a completed build job, PBF parts can have internal voids (i.e., porosity) and residual stress, which requires post-processing via stress relief, heat treatment, annealing, shot peening, hot isostatic pressing (HIP), or tomographic inspection in order to meet performance requirements [7].Figure 1.

Powder Bed Fusion Schematic (left) and Fusing of Powder Particles (right) [2]With differences pertaining to their energy source, control mechanisms, and operatingconditions [8], PBF technologies share a common feedstock in metal powder, a material which can range in price from “$260 to $450 per kg” [9]. Due to this expense, un-melted feedstock is reused, also referred to as recycled, in subsequent build jobs to save costs over purchasing additional virgin powder [10]. Despite these savings, reused powder endures partial sintering with each AM build due to latent heat from the melt-pool, leading to subsequent changes in the powder size distribution [11]. This phenomenon creates non-spherical particles and satellites leading to porosity and rough surfaces [12]. In parallel, the formation of oxides, soot, and exposure to ambient atmosphere can disseminate chemical impurities in the feedstock [7, 13].

In parallel, the formation of oxides, soot, and exposure to ambient atmosphere can disseminate chemical impurities in the feedstock [7, 13]. With many of these AM technologies having “black box” controllers [14], users are restricted to proprietary materials and process parameters provided by the Original Equipment Manufacturer (OEM), limiting options for customized processing conditions or alternative feedstock suppliers.The objective in this paper is to answer the following research question: “How does thereusability of a metal powder feedstock impact the total costs in PBF?” To answer this question, we extend an existing PBF cost model and present a case study detailing the costs for two parts built in a PBF process. Beginning with a literature review in Section 2, previous work on AM cost modeling is summarized along with current research on powder reusability in AM. Section 3 discusses the expanded cost model along with the proposed formulas and variables.

Section 3 discusses the expanded cost model along with the proposed formulas and variables. Section 4 provides an overview of the case study, followed by results and closing remarks in Section 5 and 6, respectively.2. Literature Review2.1. Cost Modeling in Additive ManufacturingLindemann et al. [15] performed a lifecycle study for production in metal additivemanufacturing (AM) using Time-driven Activity-Based Costing (ABC). Costs for a generic2008metal AM machine were discretized into an activity workflow of CAD Preparation, Machine Preparation, Build Process, Support Removal, and Surface Treatment. Considering a single part geometry for production at 4500 h/year, their cost model found that 74% of total costs can be attributed to machine costs, followed by material costs at 12%. Sensitivity analysis revealed that the material costs could vary between 5% and 46% of the total cost based on the value of the material feedstock.Rickenbacher et al.

Sensitivity analysis revealed that the material costs could vary between 5% and 46% of the total cost based on the value of the material feedstock.Rickenbacher et al. [16] expanded the work of Lindemann et al. by proposing a generic cost model for SLM. Rickenbacher et al. used ABC while accounting for multiple geometries and additional part quantities in the same AM build. Costs associated with the build time for a multiple geometry build job were evenly divided among all the parts in a layer-wise manner, based on their respective build heights. In a case-study for three geometries, Rickenbacher et al. found that a cost savings of 41% can be achieved by building up the parts together in the same build job instead of individual builds, agreeing with observations found by Lindemann et al. The total build time was identified as the largest cost driver in the workflow due to machine costs associated with exposing each layer of powder.

The total build time was identified as the largest cost driver in the workflow due to machine costs associated with exposing each layer of powder. One limitation for both of these models is that they treated the material feedstock as a fixed cost and did not allocate a cost for changes in the powder’s properties and overall value as the material was being processed and reused in PBF.2.2. Metal Feedstocks in Powder Bed FusionMetal powder is defined as a substance containing particles of elemental metals or alloys,normally less than 1000 microns in size [17]. This substance can be produced through various electro-chemical and thermo-mechanical processes such as Gas Atomization, Water Atomization, Centrifugal Disintegration [18-20]. In metallurgical applications, the powder can be characterized by the Density, Powder Size Distribution (PSD), Chemical Composition, Surface Chemistry, Morphology, Crystalline Phases, Flowability, and Thermal Properties [13, 14].

Variability in powder properties can occur at numerous stages throughout the lifecycle of the material. Axelsson [21] conducted a study on Ti-6Al-4V powder produced from three independent manufacturers for EBM. Analysis of the chemical compositions found Nitrogen, Chlorine, and Yttrium outside ASTM F2924 limits, indicating contamination during the feedstock production process. Powder is also sensitive to oxidation, a naturally occurring chemical reaction where oxygen atoms undergo diffusion and exchange electrons with a metallic element to form oxides [22]. Oxides form surface films which can alter the absorptivity and melting of the material [23].

Oxides form surface films which can alter the absorptivity and melting of the material [23]. In addition to oxidation, the build environment of PBF machines can promote the formation of carbides and nitrides due to prolonged durations and reactivity at elevated temperatures [24].Reusability refers to the ability of a powder to be reused in an AM process before it no longer conforms to standard specifications or fails to produce acceptable quality as defined by the “component supplier or purchaser” [25-28]. While reused powder can be physically sieved by particle diameter, the presence of chemical impurities, non-spherical agglomerates, and oxides may not be explicitly removed from the powder lot. In one study, Seyda et al. [29] took virgin Ti-6Al-4V powder and reused it in 12 DMLS build cycles to study changes in mechanical properties.

In one study, Seyda et al. [29] took virgin Ti-6Al-4V powder and reused it in 12 DMLS build cycles to study changes in mechanical properties. Reusing the powder caused the PSD to increase and particles to become coarse due to the formation of partially-sintered agglomerates, shown in Figure 2. This correlated to surface roughness increasing from 92 to 123 Ra after the 12 build cycles. With no post-processing, test2009specimens showed the ultimate tensile strength (UTS) increased from 1030 MPa to 1101 MPa after 6 build cycles, before decreasing to 1072 MPa for the last 6 reuses. The total amount of porosity in the produced parts decreased from 0.11% to 0.05%; however, the size of individual pores increased due to the enlarged PSD. Seyda et al. identified these pores as the cause for stress concentrations and localized defects.Figure 2: Agglomerates in Reused Ti-6Al-4V[29]Figure 3: PSD of 17-4SS [13]Slotwinski et al.

Seyda et al. identified these pores as the cause for stress concentrations and localized defects.Figure 2: Agglomerates in Reused Ti-6Al-4V[29]Figure 3: PSD of 17-4SS [13]Slotwinski et al. [13] studied stainless steel (17-4 SS) and cobalt chrome (CoCr) virginpowder, and reused them for eight DMLS build cycles to analyze changes in physical and chemical properties. Reusing the powder caused the volume fraction to increase with reuses as shown in Figure 3. Agglomerates formed in the AM process continually passed through the sieve due to their morphology or they were scrapped away by the recoating mechanism in the AM process. Overall, these particles could not be explicitly removed from the powder. The 17-4 SS and CoCor powder feedstocks were analyzed by x-ray photospectroscopy (XPS) and showed no significant difference in the elemental composition of the sieved powder after 8 build reuses.

The 17-4 SS and CoCor powder feedstocks were analyzed by x-ray photospectroscopy (XPS) and showed no significant difference in the elemental composition of the sieved powder after 8 build reuses. However, a comparison of the sieved 17-4SS powder and sieve residue revealed significant differences in the surface chemistry and crystalline phases.Tang et al. [30] reused Ti-6Al-4V ELI virgin powder for 21 build cycles in EBM, sievingthe powder at 177 microns. Oxygen content increased with each reuse from 0.08 wt.% to 0.19 wt.%, as presented in Figure 4. The oxygen content directly correlated to the UTS but not the elongation. UTS increased from 920 MPa to 1039 MPa, while elongation was scattered between 13-18%. Tang et al. identified that oxygen may be reacting with the powder as parts are being removed from the build chamber, sieving, or as parts are being cleaned in the Arcam Powder Recovery System.

Tang et al. identified that oxygen may be reacting with the powder as parts are being removed from the build chamber, sieving, or as parts are being cleaned in the Arcam Powder Recovery System. In conjunction, oxygen pick-up was attributed to exposure time in the build job along with ambient humidity in the laboratory. Since the powder progressively gained oxygen, Tang et al. suggest that Ti-6Al-4V ELI should not be reused more than 4 times in EBM to maintain compliance with ASTM F3001 (Grade 23). Alternatively, another recommendation was to replenish the reused powder lot with additional virgin powder to stabilize the oxygen content and stay “in-spec” with the provided standard.Grainger [31] conducted a similar study LM wherein Ti-6Al-4V ELI powder lot was usedin 38 build cycles, with no additional virgin powder.

The results shown in Figure 5 illustrated a2010similar trend of increasing oxygen content with each build job, and exceeding the specification for ASTM F3001 between 15-35 build cycles, yet still within acceptable limits for Grade 5 titanium. With no significant variation observed in the tensile properties and chemical pickup, Grainger concluded that there was no requirement to dispose of un-melted Ti-6Al-4V powder after it had been cycled through a number of reuses. However, Grainger noted that the results could not be extended to other AM machines and materials due to differences in how a PBF technology generates the inert atmosphere and their mechanisms for powder handling.Figure 4: Oxygen buildup in Ti-6Al-4V ELI during EBM [30]Figure 5: Oxygen buildup in Ti-6Al-4V ELI during LM [31]2.3.

Summary of Literature Review and Current GapsWith material costs ranging between 5-46% [15] of the total costs in metallic PBF processes, current cost models are assuming a constant, fixed cost, for the material feedstock and financially valuing reused powder as virgin powder. These models lack an analytical method for determining the value of the reused material as it undergoes physical and chemical changes in subsequent build jobs. The reusability of powder feedstock can vary across material alloys and PBF technologies due to the production process of the powder, energy-material interaction for a given machine, powder sieving/handling, and build chamber environments. While reactive metals such as 17-4 SS and Ti-6Al-4V ELI may attain chemical impurities with subsequent build jobs, other alloys such as IN718 may have virtually identical material composition when replenishing the lot with virgin powder [32].

While some studies have found little to no variation in the tensile properties, present literature is limited on the impact of reused powder on fatigue properties and correlating them to mixing virgin/reused feedstocks along with process parameters (e.g., power, scan velocity, hatch spacing, tool path) [23].3. Powder Costing Methodology3.1. Financial Depreciation for Costing Powder FeedstocksAs discussed in the previous section, the reusability of a powder feedstock is notexplicitly captured in current AM cost models due to the use of a fixed material cost. This creates uncertainty as to how one should financially value the feedstock as it is reused in PBF and how to precisely cost parts produced from a reused feedstock. To correlate reusability to the2011price of a powder feedstock, we propose that the material used in PBF be valued through a financial depreciation model.

To correlate reusability to the2011price of a powder feedstock, we propose that the material used in PBF be valued through a financial depreciation model. In accounting practice, depreciation is defined as the “gradual decline in the financial value of property due to increasing age and eventual obsolescence” [33]. With the risk of oxides, porosity, surface roughness, and chemical contamination, the quality of the powder feedstock diminishes as it is continually reused in PBF. Using a depreciation model, we can model the systematic loss in a feedstock’s value proportional to the powder’s degraded properties over a given duration of time.Although depreciation is traditionally used in business accounting for the United States’ Internal Revenue System [34], we are proposing depreciation strictly in the context of a costing method for PBF.

Depreciation is a function of the maximum allowable time duration for the feedstock and its salvage value, estimated market value, when it has reached the end of its useful life. Due to AM parts having individually-tailored functions and design applications, current material standards have resorted to agreement between component supplier and the purchaser for determining the acceptance of a reused powder in AM [25-28]. As discussed previously in Section 2.3, with each powder feedstock having a unique elemental composition and PBF technology, its corresponding time duration for reusability will vary based on factors related to the build chamber environment, energy-material interaction, and powder handling.Three common depreciation models are: Straight-Line (SLN), Double Declining Balance(DDB), and Sum-of-the-Year’s Digits (SOYD) [35] as shown in Figure 6.

SLN assumes a uniform reduction in value with each increment in time; however, this linear depreciation is fixed and assumes that the powder feedstock loses uniform amounts of value regardless of being at the beginning or end of its useful life. With virgin powder being at most risk of chemical contamination and oxides, an ideal depreciation model is one that reflects a relatively large rate of decline after initial uses in a PBF process and then a smaller rate as the feedstock is continually reused to correspond to compounded degradation over time. DDB presents a more accelerated model where the feedstock rapidly loses value at early stages of its useful life and then gradually less in later life; however, since DDB applies a constant multiplier for depreciation, the salvage value is not explicitly designated and therefore unadaptable to whichever value the user designates for the end-use scrap, unless manually corrected.

Serving as a median between SLN and DDB, SOYD exhibits a moderate drop in value at early life and less as the material is increasingly reused. A possible drawback for SOYD is that it does not depreciate as rapidly as the DDB. Unlike DDB, the salvage value of SOYD can be specified by the user making it a more flexible and adaptable model for a wide range of materials. Given these customizable features and appropriate depreciation rates, SOYD was selected as the most suitable model for valuing a powder feedstock as it is reused in PBF.With SOYD as the selected depreciation model, units needed to be specified for the time duration in which the powder feedstock is reused. Build cycles are widely cited in the research [29-32]; however, build cycles are an imprecise measure due to variations in the underlying build time, build height, part orientation, and quantity of parts.

While an AM build job is typically measured in units of hours [2], this value can be convoluted with the idle time and recoating time of the AM machine. While time durations can measure temporal usage, it does not provide insight on the explicit process parameters and build chamber conditions in which a powder feedstock was processed.2012Figure 6: Comparison of Depreciation Methods for Powder FeedstockGiven the present information in literature and limited alternatives, build cycles was selected as the unit for measuring the time duration of a powder feedstock in PBF. Equation 1 is our proposed method for valuing the powder feedstock using SOYD. For this equation, it is assumed that a powder lot is reused on a single AM machine, where the overflow, part bed, and feed bed powder are mixed and sieved after each build.

For this equation, it is assumed that a powder lot is reused on a single AM machine, where the overflow, part bed, and feed bed powder are mixed and sieved after each build. Builds are conducted using consistent process parameters, atmosphere, material handling conditions, and have no added virgin powder.where:𝐶𝐶𝑚𝑚𝑢𝑢+1 = 𝐶𝐶𝐶𝐶𝑢𝑢 − (𝐶𝐶𝐶𝐶0 − 𝑆𝑆) ∙ (𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚 − 𝑢𝑢 + 1 𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚(𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚 + 1) 2)is the cost of the powder feedstock that has been used u times ($/kg), is the cost of a virgin powder feedstock ($/kg),𝐶𝐶𝐶𝐶𝑢𝑢 is the salvage value of the powder at the end of its depreciable life ($/kg), 𝐶𝐶𝐶𝐶0 𝑆𝑆 is the number of build cycles a powder has underwent in PBF (-). 𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚 𝑢𝑢is the maximum number of build cycles a powder can be used for a PBF technology (-),3.2. Activity-Based Cost Modeling for Powder Bed FusionTo implement the proposed model for valuing a reused powder, we expand upon the work of Rickenbacher et al.

Activity-Based Cost Modeling for Powder Bed FusionTo implement the proposed model for valuing a reused powder, we expand upon the work of Rickenbacher et al. [16] and follow a similar workflow to that shown in Figure 7. A built-up AM part, Pi, shall consist of geometry, Gi with Ni quantity in a build job for PBF. Due to overlap with the previous model by Rickenbacher et al., we only present equations that have been modified or are unique to this paper.2013(1)Figure 7: Rickenbacher et al.’s Workflow for SLM [16]The first labor activity in this cost model is the preparation of the digital geometry data[16]. Upon receiving the costumer’s digital model, it is assumed that the geometry meets general Design for AM rules (e.g., process selection, fully enclosed surfaces, wall thickness, tolerances) [36-38]. The tasks in this activity are selecting a build orientation and generating support structures.

The tasks in this activity are selecting a build orientation and generating support structures. Software packages such as NetFabb and Materialise Magics [39, 40] can automate these processes; however, these programs are not robust for metal AM and may produce designs that satisfy manufacturing requirements but fail to meet product specifications [41]. Thus, this activity is an iterative process that relies on experiential knowledge of the AM operator and must be tailored to the given PBF technology and geometry.Once all of the parts have been successfully prepared for AM, the digital geometries are read into a build layout program. The manipulation and placement of geometries on the digital build tray can have repercussions regarding the likelihood of build failure (e.g., collision with recoating mechanism, curling, surface roughness) [6, 42].

With limited literature and standards for this activity, it is also conducted in an iterative manner relying on previous knowledge and past experience. The time required for arranging the geometries is a function of the total part geometries and replicates present in the AM build job.Machine set-up consists of uploading the digital build tray files, selecting processparameters, initializing inert gas (or vacuum depending on PBF technology), and readying system hardware. With metal powders having explosive and physiological hazards [2, 7, 43-45], material handling is dangerous, requiring timely and duteous tasks for safe activity. Additional time can occur if the build calls for a different material than the one currently loaded in the machine.

Additional time can occur if the build calls for a different material than the one currently loaded in the machine. The total time for changing materials includes the tasks of unloading the current powder, cleaning the build chamber, replacing consumables (e.g., filters, inert gas), loading the new feedstock, and cleaning all ancillary equipment (i.e., vacuum). While Rickenbacher et al.

For build jobs with multiple parts, this formula assumes that the build rate for exposing each voxel is constant and that all time latency due to positioning the energy scanner between consolidated powder regions is negligible. Using the algorithm proposed by Rickenbacher et al.

Using the algorithm proposed by Rickenbacher et al. [16], the recoating time for build jobs with multiple build heights is calculated in a time fraction manner for each part:is the AM machine’s hourly rate ($/h), is the operator’s hourly rate ($/hour), is the time for setting up the machine (h),is the total time for changing the material, including corresponding activities (h),is the quantity of parts with ith geometry (-).The derivation of a high-fidelity build-time estimator is outside the scope of this paper.+(3)𝑇𝑇𝑖𝑖𝑖𝑖𝑖𝑖𝑒𝑒 𝑖𝑖 ∑ 𝑁𝑁𝑖𝑖+ 𝑇𝑇𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏 𝑠𝑠𝑠𝑠𝑠𝑠𝑠𝑠𝑏𝑏 ∙ ∑ �𝑁𝑁𝑖𝑖 ∙ 𝑉𝑉𝑠𝑠𝑜𝑜𝑠𝑠𝑚𝑚𝑏𝑏𝑖𝑖� is the total time required for building up a single part in a given build job (h),𝑇𝑇𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏(𝑃𝑃𝑖𝑖) =𝑖𝑖𝑇𝑇𝑜𝑜𝑠𝑠𝑀𝑀𝑜𝑜𝑚𝑚𝑠𝑠(𝑃𝑃𝑖𝑖)where:is the built-up AM part corresponding to ith geometry (-),𝑇𝑇𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏 𝑃𝑃𝑖𝑖 𝑇𝑇𝑖𝑖𝑏𝑏𝑏𝑏𝑠𝑠 𝑇𝑇𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏 𝑠𝑠𝑠𝑠𝑠𝑠𝑠𝑠𝑏𝑏 𝑁𝑁𝑖𝑖 𝑉𝑉𝑠𝑠𝑜𝑜𝑠𝑠𝑚𝑚𝑏𝑏𝑖𝑖 𝑇𝑇𝑜𝑜𝑠𝑠𝑀𝑀𝑜𝑜𝑚𝑚𝑠𝑠 commences, and the AM machine proceeds to fabricate the designated part(s).

To cost the activity of the AM machine throughout this duration, costs are grouped into three categories as shown in Equation 4. Machine costs includes costs pertaining to AM machine utilization and inert gas, multiplied by the build time allocated to a given part (see Equation 5). Material cost, as presented in Equation 6, is the cost related to the mass of the powder feedstock melted by the AM machine to produce the part. Using the concepts introduced in Section 3.1, the material costs is valued as a function of the build cycles endured by the material feedstock loaded in the AM machine.

For the mass of the part, in Equation 7, empirical factors are included to compensate for powder losses due to particles trapped in the filters [48] during processing and loose powder trapped in hollow support structures.is the time when the AM machine is inactive (e.g., heating, cooling) (h),is the average time for the AM machine to consolidate a voxel of powder (is the quantity of parts with ith geometry (-),ℎ/𝑐𝑐𝐶𝐶is the total volume of the part and support structures for ith geometry ( is the total recoating time allocated to a single part (h).),3𝑐𝑐𝐶𝐶Once the operator has completed all hardware and software set-up, then the build job3),where:𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏(𝑃𝑃𝑖𝑖) = 𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑀𝑀𝑚𝑚𝑀𝑀ℎ𝑖𝑖𝑎𝑎𝑠𝑠(𝑃𝑃𝑖𝑖) + 𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑀𝑀𝑚𝑚𝑠𝑠𝑠𝑠𝑜𝑜𝑖𝑖𝑚𝑚𝑏𝑏(𝑃𝑃𝑖𝑖) + 𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝐷𝐷𝑠𝑠𝑠𝑠𝑜𝑜𝑠𝑠𝑀𝑀𝑖𝑖𝑚𝑚𝑠𝑠𝑖𝑖𝑜𝑜𝑎𝑎(𝑃𝑃𝑖𝑖) (4)is the cost per part for building up a part using the AM machine ($),𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏2015is the built-up AM part corresponding to ith geometry (-),is the cost per part for operating the AM machine during a build job ($), is the cost per part for the melted powder feedstock in the AM process ($),𝑃𝑃𝑖𝑖 𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑚𝑚𝑚𝑚𝑀𝑀ℎ𝑖𝑖𝑎𝑎𝑠𝑠 𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑚𝑚𝑚𝑚𝑠𝑠𝑠𝑠𝑜𝑜𝑖𝑖𝑚𝑚𝑏𝑏 𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝐷𝐷𝑠𝑠𝑠𝑠𝑜𝑜𝑠𝑠𝑀𝑀𝑖𝑖𝑚𝑚𝑠𝑠𝑖𝑖𝑜𝑜𝑎𝑎is the cost per part for the un-melted feedstock used in the AM process ($).where:𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑚𝑚𝑚𝑚𝑀𝑀ℎ𝑖𝑖𝑎𝑎𝑠𝑠(𝑃𝑃𝑖𝑖) = 𝑇𝑇𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏 ∙ �𝐶𝐶𝑚𝑚𝑚𝑚𝑀𝑀ℎ + 𝐶𝐶𝑎𝑎𝑚𝑚𝑠𝑠�is the cost per part for producing a build job in the AM process ($),is the built-up AM part corresponding to ith geometry (-),𝐶𝐶𝑚𝑚𝑚𝑚𝑀𝑀ℎ𝑖𝑖𝑎𝑎𝑠𝑠 𝑃𝑃𝑖𝑖 𝑇𝑇𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏 𝐶𝐶𝑚𝑚𝑚𝑚𝑀𝑀ℎ 𝐶𝐶𝑎𝑎𝑚𝑚𝑠𝑠 𝑁𝑁𝑖𝑖is the time for building up the entire job in the AM process (h), is the AM machine’s hourly operating cost ($/h),is the cost for inert gas consumption during the build ($/h),is the quantity of parts with ith geometry (-).where:𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑚𝑚𝑚𝑚𝑠𝑠𝑠𝑠𝑜𝑜𝑖𝑖𝑚𝑚𝑏𝑏(𝑃𝑃𝑖𝑖) = 𝑀𝑀𝑖𝑖 ∙ 𝐶𝐶𝐶𝐶𝑢𝑢is the cost per part for the powder feedstock melted in the AM process ($),is the built-up AM part corresponding to ith geometry (-), is the mass of a part with ith geometry (kg),𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑚𝑚𝑚𝑚𝑠𝑠𝑠𝑠𝑜𝑜𝑖𝑖𝑚𝑚𝑏𝑏 𝑃𝑃𝑖𝑖 𝑀𝑀𝑖𝑖 𝐶𝐶𝐶𝐶𝑢𝑢is the cost of the powder feedstock that has been used in u build cycles ($/kg).where:𝑀𝑀𝑖𝑖 = (1 + 𝛼𝛼) ∙ 𝜌𝜌𝑤𝑤 ∙ (𝑉𝑉𝑠𝑠𝑚𝑚𝑜𝑜𝑠𝑠𝑖𝑖 + 𝑉𝑉𝑠𝑠𝑢𝑢𝑠𝑠𝑠𝑠𝑜𝑜𝑜𝑜𝑠𝑠𝑠𝑠𝑖𝑖 ) + 𝛾𝛾 ∙ 𝜌𝜌𝑠𝑠 ∙ 𝑉𝑉𝑠𝑠𝑢𝑢𝑠𝑠𝑠𝑠𝑜𝑜𝑜𝑜𝑠𝑠𝑠𝑠𝑖𝑖is the mass of a part with ith geometry (kg),is the percentage of powder loss due to process inefficiency (%), 𝑀𝑀𝑖𝑖 is the percentage of powder loss due to being trapped within support structures (%), 𝛼𝛼 𝛾𝛾 𝑉𝑉𝑠𝑠𝑚𝑚𝑜𝑜𝑠𝑠𝑖𝑖 𝑉𝑉𝑠𝑠𝑢𝑢𝑠𝑠𝑠𝑠𝑜𝑜𝑜𝑜𝑠𝑠𝑠𝑠𝑖𝑖 𝜌𝜌𝑤𝑤 𝜌𝜌𝑠𝑠 In the third cost category, a build job requires powder feedstock to fill the powder bed and support built-up geometries throughout processing.

This un-melted powder can become degraded by soot, oxides, and agglomerates during the AM process. These byproducts diminish the financial value of the feedstock, regardless of subsequent sieving, because the un-melted powder becomes populated with impurities that can propagate into future layers or builds. Thus, the un-melted powder loses the opportunity to be implemented as a virgin powder and produce parts with minimal defects from the base material. To allocate cost for this phenomenon, we propose that the financial value lost by the surrounding un-melted powder be charged to all parts produced within the build job. For this third category, the proposed cost is defined as Depreciation and calculated using Equation 8:is the volume of the part body for the ith geometry (c),is the volume of the support structures for the ith geometry (c),3𝐶𝐶is the powder wrought density ( is the powder tap density (),3𝐶𝐶).

𝑘𝑘𝑘𝑘/𝑐𝑐𝐶𝐶 33𝑘𝑘𝑘𝑘/𝑐𝑐𝐶𝐶2016(5)(6)(7)𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝐷𝐷𝑠𝑠𝑠𝑠𝑜𝑜𝑠𝑠𝑀𝑀𝑖𝑖𝑚𝑚𝑠𝑠𝑖𝑖𝑜𝑜𝑎𝑎(𝑃𝑃𝑖𝑖) =where:𝑀𝑀𝑖𝑖 ∑ (𝑁𝑁𝑖𝑖 ∙ 𝑀𝑀𝑖𝑖)𝑖𝑖�𝑀𝑀𝐹𝐹𝐹𝐹 − � (𝑁𝑁𝑖𝑖 ∙ 𝑀𝑀𝑖𝑖) 𝑖𝑖� ∙ (𝐶𝐶𝐶𝐶𝑢𝑢 − 𝐶𝐶𝐶𝐶𝑢𝑢+1)is the cost per part for the un-melted feedstock used in the AM process ($),is the built-up AM part corresponding to ith geometry (-), is the mass of a part with ith geometry (kg), is the quantity of parts with ith geometry (-),𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝐷𝐷𝑠𝑠𝑠𝑠𝑜𝑜𝑠𝑠𝑀𝑀𝑖𝑖𝑚𝑚𝑠𝑠𝑖𝑖𝑜𝑜𝑎𝑎 𝑃𝑃𝑖𝑖 𝑀𝑀𝑖𝑖 𝑁𝑁𝑖𝑖 𝑀𝑀𝐹𝐹𝐹𝐹 𝐶𝐶𝐶𝐶𝑢𝑢 the feed bed and subtracting the total mass of all built-up parts, including their corresponding powder losses. This is multiplied by the difference in financial value of the feedstock at its present state to the diminished value after one additional build cycle. Parts within the build job are allocated the Depreciation cost as a function of their mass fractions relative to the total mass of all built-up parts.

Parts within the build job are allocated the Depreciation cost as a function of their mass fractions relative to the total mass of all built-up parts. Through Depreciation, this costing method accounts for the melting and, in- parallel, the degradation of the un-melted powder feedstock when building up a part in PBF. The calculation of the powder mass loaded in the feed bed is non-trivial and must be sufficient to fill the part bed completely and build-up the geometries.

The calculation of the powder mass loaded in the feed bed is non-trivial and must be sufficient to fill the part bed completely and build-up the geometries. The amount of loaded powder feedstock is influenced by the material type and how a PBF technology accounts for changes in the levelling [7] of a layer as consolidated powder regions melt and re-solidify.is the total mass of the powder loaded into the AM machine’s feed bed (kg), is the cost of the powder feedstock that has been used in u build cycles ($/kg).In Equation 8, Depreciation cost is calculated by taking the mass of the powder loaded inOnce the AM process has finished building, the build job and all subsequent parts arephysically removed from the AM machine. Equation 9 [16] takes the time required for this task and evenly divides it among the total number of parts created in the build job.

Equation 9 [16] takes the time required for this task and evenly divides it among the total number of parts created in the build job. Activities at this stage pertain to removing the build substrate from the machine, collecting all loose un-melted powder from the part bed, cleaning the machine, removing all powder from the feed bed and overflow bins, sieving the used powder, storage, and documentation. Empirical factors for extra effort under inert environment have been removed from the original model.where:𝐶𝐶𝑅𝑅𝑠𝑠𝑚𝑚𝑜𝑜𝑅𝑅𝑚𝑚𝑏𝑏(𝑃𝑃𝑖𝑖) = (𝐶𝐶𝑜𝑜𝑠𝑠 + 𝐶𝐶𝑚𝑚𝑚𝑚𝑀𝑀ℎ) ∙𝑇𝑇𝑅𝑅𝑠𝑠𝑚𝑚 𝑖𝑖 ∑ 𝑁𝑁𝑖𝑖is the cost per part for removing the substrate/parts from the AM machine ($),is the built-up AM part corresponding to ith geometry (-),𝐶𝐶𝑅𝑅𝑠𝑠𝑚𝑚𝑜𝑜𝑅𝑅𝑚𝑚𝑏𝑏 𝑃𝑃𝑖𝑖 𝑇𝑇𝑅𝑅𝑠𝑠𝑚𝑚 𝐶𝐶𝑜𝑜𝑠𝑠𝑠𝑠𝑜𝑜 𝐶𝐶𝑀𝑀𝑚𝑚𝑀𝑀ℎ 𝑁𝑁𝑖𝑖 The next step involves separation of parts from the build substrate.

Modifying the original formula [16], parts produced in a build job using laser-powered PBF may have built-up residual stress and thus undergo a stress-relief [49] to reduce geometric distortion upon separation. Once completed, wire electrical discharge machining (EDM) is used to physicallyis the time required to remove parts, clean machine, and perform all ancillary tasks (h), is the operator’s hourly rate ($/hour), is the AM machine’s hourly operating cost ($/h),is the quantity of parts with ith geometry (-).2017(8)(9)detach all parts from the substrate.

The costs for wire EDM are allocated based on the contact area occupied by a part, and their support structures, on the substrate as follows in Equation 10.where:𝐴𝐴𝑀𝑀𝑜𝑜𝑎𝑎(𝐺𝐺𝑖𝑖) 𝐶𝐶𝑠𝑠𝑠𝑠𝑜𝑜𝑠𝑠𝑠𝑠𝑠𝑠 ∑ 𝑁𝑁𝑖𝑖 ∙ 𝐴𝐴𝑀𝑀𝑜𝑜𝑎𝑎(𝐺𝐺𝑖𝑖) 𝑖𝑖 ∑ 𝑁𝑁𝑖𝑖 is the cost per part for separating a part from the substrate ($),𝐶𝐶𝑆𝑆𝑢𝑢𝑏𝑏𝑠𝑠𝑠𝑠𝑜𝑜𝑚𝑚𝑠𝑠𝑠𝑠(𝑃𝑃𝑖𝑖) =+ 𝐶𝐶𝐸𝐸𝐷𝐷𝑀𝑀 ∙𝑖𝑖is the built-up AM part corresponding to ith geometry (-), is the cost for stress-relieving a build plate($), is the total cost for separating a part via EDM ($), is the connected area of a part to the substrate (𝐶𝐶𝑆𝑆𝑢𝑢𝑏𝑏𝑠𝑠𝑠𝑠𝑜𝑜𝑚𝑚𝑠𝑠𝑠𝑠 𝑃𝑃𝑖𝑖 𝐶𝐶𝑠𝑠𝑠𝑠𝑜𝑜𝑠𝑠𝑠𝑠𝑠𝑠 𝐶𝐶𝐸𝐸𝐷𝐷𝑀𝑀 𝐴𝐴𝑀𝑀𝑜𝑜𝑎𝑎 𝑁𝑁𝑖𝑖 undergo additional post-processing to meet customer requirements. Due to parts having individually-tailored functions and applications, the required operations and sequence of their events will vary due to the specifications ordered by the costumer. For Equation 11 [16], cost is calculated for post-processing based on support structure removal for an individual part.

Costs are allocated based on an AM operator’s labor for part preparation, arranging geometries on the build tray, setting-up the AM machine, executing the build job, and removing the substrate and decommissioning the machine. Afterwards, post-processing of the build job begins with stress-relief, wire EDM, followed last by individual post-processing to produce a fully- functional component.

These 7 activities are added together in Equation 12 to produce the total cost for a part made using PBF:is the time required to post-process a part geometry ($),is the ith geometry (-),is the operator’s hourly rate ($/hour),is the hourly rate of tools and machines for post-processing ($).𝐶𝐶𝑇𝑇𝑜𝑜𝑠𝑠𝑚𝑚𝑏𝑏(𝑃𝑃𝑖𝑖) = 𝐶𝐶𝑠𝑠𝑜𝑜𝑠𝑠𝑠𝑠(𝑃𝑃𝑖𝑖) + 𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏𝑏𝑏𝑜𝑜𝑏𝑏(𝑃𝑃𝑖𝑖) + 𝐶𝐶𝑆𝑆𝑠𝑠𝑠𝑠𝑢𝑢𝑠𝑠(𝑃𝑃𝑖𝑖) + 𝐶𝐶𝐹𝐹𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏(𝑃𝑃𝑖𝑖) + 𝐶𝐶𝑅𝑅𝑠𝑠𝑚𝑚𝑜𝑜𝑅𝑅𝑚𝑚𝑏𝑏(𝑃𝑃𝑖𝑖)+ 𝐶𝐶𝑆𝑆𝑢𝑢𝑏𝑏𝑠𝑠𝑠𝑠𝑜𝑜𝑚𝑚𝑠𝑠𝑠𝑠(𝑃𝑃𝑖𝑖) + 𝐶𝐶𝑃𝑃𝑜𝑜𝑠𝑠𝑠𝑠𝑠𝑠(𝑃𝑃𝑖𝑖)2018(10)(11)(12)where:is the total manufacturing costs ($), is the cost per part for preparing the digital geometry data ($),𝐶𝐶𝑠𝑠𝑜𝑜𝑠𝑠𝑚𝑚𝑏𝑏 𝐶𝐶𝑠𝑠𝑜𝑜𝑠𝑠𝑠𝑠 𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏𝑏𝑏𝑜𝑜𝑏𝑏 𝐶𝐶𝑠𝑠𝑠𝑠𝑠𝑠𝑢𝑢𝑠𝑠 𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏 𝐶𝐶𝑅𝑅𝑠𝑠𝑚𝑚𝑜𝑜𝑅𝑅𝑚𝑚𝑏𝑏 𝐶𝐶𝑆𝑆𝑢𝑢𝑏𝑏𝑠𝑠𝑠𝑠𝑜𝑜𝑚𝑚𝑠𝑠𝑠𝑠 𝐶𝐶𝑠𝑠𝑜𝑜𝑠𝑠𝑠𝑠𝑠𝑠is the cost per part for the build tray assembly ($),is the cost for setting up the machine ($), is the cost for building up a part in the AM process ($),is the cost for removing the substrate/parts from the machine ($), is the cost for separating a part from the substrate ($),is the total cost for post-processing ($).4.

Implementation and Case StudyTo demonstrate this approach, costs were studied for two parts designed for and producedby laser-based PBF. Figures 8 and 9 show the geometric data and build orientation for each of the parts, which were built using EOS Titanium Ti64 and Stainless Steel GP1, respectively. These parts were selected because Ti64 is a high-value material and GP1 is relatively cheaper. All parts were manufactured on an EOSINT M280 DMLS machine at Penn State’s Center for Innovative Materials Processing through Direct Digital Deposition (CIMP-3D). Following the workflow of Rickenbacher et al., STL files were first imported into Materialise Magics for digital preparation and support structure generation. Geometries were then entered into EOS RP Tools and sliced to form the build job file. Next, the files were entered into EOS PSW, where process parameters were selected, and then the AM process commenced.

Next, the files were entered into EOS PSW, where process parameters were selected, and then the AM process commenced. Once completed, the build substrate was removed from the machine, the machine was cleaned, powder was sieved, and build documentation was completed. Finally, the parts and substrate were shipped to a local manufacturer for stress-relief and wire EDM, before returning to CIMP-3D for post-processing.Figure 8: Automotive Upright – Geometry G1Figure 9: Testing Apparatus– Geometry G2All equations from Section 3 were written in a MATLAB program and used for costing each of the builds. For this case study, constants for powder losses due to process inefficiency ( based on CIMP-3D staff’s experience from three years of producing metal AM parts at the 𝛼𝛼) 𝛾𝛾) facility [48].

For this case study, constants for powder losses due to process inefficiency ( based on CIMP-3D staff’s experience from three years of producing metal AM parts at the 𝛼𝛼) 𝛾𝛾) facility [48]. Unique to laser-powered PBF is that the part bed lowers by a layer thickness andand powder trapped in support structures (were chosen to be 40% and 25%, respectively,2019the feed bed dispenser platform rises between “2x or 2.5x” [7] the layer thickness to account for changes in powder levelling during thermal consolidation. The ratio of the vertical rise of the feed bed to the lowering of the part bed shall be referred to as CA, also known as charge [50]. If CA is set too low by the user, then the build can be prone to powder shorting and insufficient coverage of the build plate [7, 50]. Thus, with the guidance of the CIMP-3D staff, a CA of 2.25 was assumed to be sufficient for all build jobs in this case study.

Thus, with the guidance of the CIMP-3D staff, a CA of 2.25 was assumed to be sufficient for all build jobs in this case study. Equation 13 was used to estimate the mass of powder loaded in the feed bed for a generic laser-powered PBF machine. A list of constants used in the cost model for these two examples are included in Table 1.(13)where:𝑀𝑀𝐹𝐹𝐹𝐹 = 𝐶𝐶𝐴𝐴 ∙ 𝐷𝐷𝑚𝑚 ∙ 𝐷𝐷𝑦𝑦 ∙ 𝐵𝐵ℎ(𝑃𝑃𝑖𝑖) ∙ 𝜌𝜌𝑠𝑠is the total mass of the powder loaded into the AM machine’s feed bed (kg),is the vertical rise of the feed bed per layer thickness in the build (-), is the length of the dispenser platform in the feed bed (mm), is the width of the dispenser platform in the feed bed (mm), is the build height of the tallest part in the build job (mm), is the powder tap density (𝑀𝑀𝐹𝐹𝐹𝐹 𝐶𝐶𝐴𝐴 𝐷𝐷𝑚𝑚 𝐷𝐷𝑦𝑦 𝐵𝐵ℎ 𝜌𝜌𝑠𝑠 Table 2 contains all the values used in the cost model for the two example parts. These values were collected from material datasheets available from the manufacturer [51, 52].

These values were collected from material datasheets available from the manufacturer [51, 52]. Since the feedstocks have a wide range of permissible reuses, costs were modeled parametrically by , between 10, 20, 30, and 40, to correspond varying the maximum allowable build cycles, with the data found in the literature review. The graph in Figure 10 shows an example of the 𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚 cost for a powder feedstock at each of the maximum build cycles. Each point represents the financial value of a powder feedstock with the given number of accumulated reuses. The salvage value, estimated resale value, of a powder that exceeded the maximum amount of build cycles was assumed to be zero.).3𝑘𝑘𝑘𝑘/𝑐𝑐𝐶𝐶Figure 10: Powder Feedstock Value vs. Build Cycles vs.

Maximum Build Cycles2020Table 1: Cost Model ConstantsVariable𝐶𝐶𝑜𝑜𝑠𝑠𝑠𝑠𝑜𝑜 𝐶𝐶𝑠𝑠𝑀𝑀 𝐶𝐶𝑀𝑀𝑚𝑚𝑀𝑀ℎ 𝐶𝐶𝑎𝑎𝑚𝑚𝑠𝑠 𝑇𝑇𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏𝑏𝑏𝑜𝑜𝑏𝑏 𝑇𝑇𝑠𝑠𝑠𝑠𝑠𝑠𝑢𝑢𝑠𝑠 𝑇𝑇𝑚𝑚𝑚𝑚𝑠𝑠.𝑀𝑀ℎ𝑚𝑚𝑎𝑎𝑎𝑎𝑠𝑠 𝑇𝑇𝑖𝑖𝑏𝑏𝑏𝑏𝑠𝑠 𝑇𝑇𝑜𝑜𝑠𝑠𝑀𝑀𝑜𝑜𝑚𝑚𝑠𝑠 𝑜𝑜𝑚𝑚𝑠𝑠𝑠𝑠 𝑇𝑇𝑜𝑜𝑠𝑠𝑚𝑚 𝛼𝛼 𝛾𝛾 𝑇𝑇𝑅𝑅𝑠𝑠𝑚𝑚 𝐶𝐶𝑠𝑠𝑠𝑠𝑜𝑜𝑠𝑠𝑠𝑠𝑠𝑠 𝐶𝐶𝐸𝐸𝐷𝐷𝑀𝑀 𝐶𝐶𝑠𝑠𝑜𝑜𝑜𝑜𝑏𝑏𝑠𝑠 𝐶𝐶𝐴𝐴 𝐷𝐷𝑚𝑚 𝐷𝐷𝑦𝑦 𝑆𝑆𝑆𝑆𝑚𝑚 𝑆𝑆𝑆𝑆𝑦𝑦Description Operator’s hourly rate Cost for using computer workstation with all software and licenses AM machine’s hourly rate Cost for inert gas consumption during the build Time required for arranging all geometries in the build job Time required to set up the AM machineTime for changing and loading new powder into AM machineTime when AM machine is inactive in build (startup, heating, cooling)Average time for AM machine to spread one layer of powderTime required to remove substrate and clean machine after build Percentage of powder loss due to process inefficiency Percentage of powder loss due to entrapment in support structures Time required to remove parts, clean machine after build Total cost for thermally processing all parts on build substrate Total cost for separating all parts on substrate via EDM Cost for work area with tools and machines for post-processing Vertical rise of the feed bed per layer thickness in the build Length of the dispenser platform in the feed bed Width of the dispenser platform in the feed bed Length of the build substrate platform in the part bed Width of the build substrate platform in the part bedValue Units 110 100 60 10 1 2 3 - 9 3 40 25 3 350 200 50 2.25 228 250 250 250$/h $/h $/h $/h h h h h sec h % % h $ $ $ - cm cm cm cmTable 2: Material ConstantsTi64 𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴 GP1𝝆𝝆𝑴𝑴 𝟑𝟑 2.74 (𝒈𝒈/𝒄𝒄𝒎𝒎 5.3)𝝆𝝆𝒘𝒘 𝟑𝟑 4.41 (𝒈𝒈/𝒄𝒄𝒎𝒎 7.8)𝑳𝑳𝑻𝑻 30 (𝝁𝝁𝒎𝒎 20)) 𝑻𝑻𝒃𝒃𝒃𝒃𝑴𝑴𝑴𝑴𝒃𝒃 𝒔𝒔𝒔𝒔𝑴𝑴𝑴𝑴𝒃𝒃 𝟑𝟑 13.5 (𝒉𝒉/𝒄𝒄𝒎𝒎 7.2𝑪𝑪𝒎𝒎𝟎𝟎 680 ($/𝒌𝒌𝒈𝒈 100))𝑺𝑺 - ($ -) 𝑼𝑼𝒎𝒎𝑴𝑴𝒎𝒎 {10,20,30,40} (− {10,20,30,40}5.

Results and Discussion5.1. Labor Time and Build ResultsTable 3 lists build data for the two example parts. Geometryhad a build time of 55hours, and Geometry predicted for both parts by 6%. This discrepancy is attributed to the assumption that for preheating, machine cool-down, and laser positioning between hatches during the build process. time of this study, there was insufficient records that documented the build cycles endured by the powder feedstocks used in the production of the two example parts. To accommodate for this lack of information, costing scenarios were generated for all material reuses ranging from virgin powder, reused powder, and powders that reached their maximum allowable build cycles.had a build time of 31 hours. The estimated build time was over-𝐺𝐺1is zero𝐺𝐺2𝑇𝑇𝑖𝑖𝑏𝑏𝑏𝑏𝑠𝑠required 61 kg of powder to perform the build, whereasrequired 32 kg.

The estimated build time was over-𝐺𝐺1is zero𝐺𝐺2𝑇𝑇𝑖𝑖𝑏𝑏𝑏𝑏𝑠𝑠required 61 kg of powder to perform the build, whereasrequired 32 kg. At the𝐺𝐺1𝐺𝐺22021Table 3: Labor Time and Build Results𝑷𝑷𝑴𝑴𝑴𝑴𝑴𝑴 𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴𝑴Ti64 GP1𝐺𝐺1 𝐺𝐺2) 𝑵𝑵𝑴𝑴 (− 1 1) 𝑽𝑽𝑴𝑴𝒕𝒕𝑴𝑴𝑴𝑴𝑴𝑴 𝟑𝟑 (𝒄𝒄𝒎𝒎 587 192𝑩𝑩𝒉𝒉 (𝒎𝒎𝒎𝒎 172 47)) 𝑻𝑻𝒔𝒔𝑴𝑴𝑴𝑴𝒔𝒔 3 (𝒉𝒉 2) 𝑻𝑻 𝒃𝒃𝒃𝒃𝑴𝑴𝑴𝑴𝒃𝒃 𝑴𝑴𝒔𝒔𝑴𝑴𝑴𝑴𝒎𝒎𝑴𝑴𝑴𝑴𝑴𝑴 58 (𝒉𝒉 335.2. Costing for Example Parts) 𝑻𝑻 𝒃𝒃𝒃𝒃𝑴𝑴𝑴𝑴𝒃𝒃 𝑴𝑴𝒄𝒄𝑴𝑴𝒃𝒃𝑴𝑴𝑴𝑴 55 (𝒉𝒉 31) 𝑻𝑻𝒔𝒔𝒕𝒕𝒔𝒔𝑴𝑴𝒔𝒔 3 (𝒉𝒉 2) 𝑴𝑴𝑴𝑴 (𝒌𝒌𝒈𝒈 3.9 2.2Figure 11 shows the range of total costs, including all labor activities, for Geometry a function of the build cycles endured by the feedstock loaded for the build job. The points along the graph represent the cost for a build job loaded with a powder feedstock with the given number of reuses. Powders having a having the shortest allowable reuses and therefore the most rapid decline in value.

Powders having a having the shortest allowable reuses and therefore the most rapid decline in value. Meanwhile, powders with a 40 build cycle limit had a longer reuse duration and thus a slower rate of decline.𝐺𝐺1 of 10 build cycles were the most expensive due to𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚The total cost for manufacturing these parts ranged among virgin powders, represented as zero build cycle feedstocks, between $11,000 and $16,500. The lowest cost scenario was the use of a powder that exceeded the maximum amount of permissible build cycles (i.e., powder that is chemically out-of-specification, diminished flowability, etc.), and total costs were $6800. This is because as a powder is increasingly reused, proportion to the diminishing financial value of the powder feedstock with each build cycle. At this lower limit, the feedstock’s value has been reduced to the salvage value, zero for this case study.

At this lower limit, the feedstock’s value has been reduced to the salvage value, zero for this case study. Hence, the material cost has diminished and the depreciation cost has become zero, due to a zero difference in the financial value of a feedstock that has become a scrap material.anddecrease in𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑚𝑚𝑚𝑚𝑠𝑠𝑠𝑠𝑜𝑜𝑖𝑖𝑚𝑚𝑏𝑏𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑏𝑏𝑠𝑠𝑠𝑠𝑜𝑜𝑠𝑠𝑀𝑀𝑖𝑖𝑚𝑚𝑠𝑠𝑖𝑖𝑜𝑜𝑎𝑎Figure 11: Total Costs vs. Build Cycles for PowderFigure 12: Cost for Workflow Activities forReuse for GeometryGeometrywith Virgin Powder𝐺𝐺1𝐺𝐺1The previous Rickenbacher et al. model, which used fixed material costs, had a constant value of $9400, regardless of the reuses accumulated by the input powder. In comparison to the cost scenarios accounting for reused powders, the fixed material cost method undervalued the cost of build jobs with virgin Ti64 powder within a range of 13% and 75%.

In comparison to the cost scenarios accounting for reused powders, the fixed material cost method undervalued the cost of build jobs with virgin Ti64 powder within a range of 13% and 75%. After 7 and/or 13 build cycles, the fixed material cost model started to overvalue total costs and thus, build jobs2022) 𝑴𝑴𝑭𝑭𝑩𝑩 (𝒌𝒌𝒈𝒈 61 32asusing a powder that had surpassed these cycles could achieve a cost savings. The largest cost savings was a 38% reduction compared to the fixed material cost method, specifically when using a powder that exceeds its useful life ().𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚Figure 12 shows a cost breakdown using virgin powder and each of their maximum ,reuses. The top three costs are depreciation cost was the largest cost for a powder with a 10 build cycle limit at 42% of the total 𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑚𝑚𝑚𝑚𝑀𝑀ℎ𝑖𝑖𝑎𝑎𝑠𝑠 cost. When machine cost, and subsequently minimized to 17% of costs when cycles. In comparison, the depreciation costs were more than 2.6 times the material costs., and.

When machine cost, and subsequently minimized to 17% of costs when cycles. In comparison, the depreciation costs were more than 2.6 times the material costs., and. The𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑚𝑚𝑚𝑚𝑠𝑠𝑠𝑠𝑜𝑜𝑖𝑖𝑚𝑚𝑏𝑏 was greater than 20 build cycles, the depreciation cost was overtaken by the was equal to 40 build𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑏𝑏𝑠𝑠𝑠𝑠𝑜𝑜𝑠𝑠𝑀𝑀𝑖𝑖𝑚𝑚𝑠𝑠𝑖𝑖𝑜𝑜𝑎𝑎𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚and displayed in Figures 13 and 14. Builds using virgin GP1 powder had a range between $5000 and $5500. The lower limit for builds using powders that exceeded reference, which valued the builds at approximately $4900, virgin GP1 powder builds were undervalued between 3% and 11%, depending on the maximum reuses permitted for the feedstock. Upon surpassing 6 and/or 11 build cycles, the totals costs became overvalued, and thus a cost savings of 5% could be achieved by building with a powder outside the allowable for at 11%,Similar analysis was conducted on Geometry𝐺𝐺2was $4600.

Using the fixed material cost model as the𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚. The largest cost for Geometrywasat 42%, , when using a 10 build cycleand 𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚 maximum virgin powder, the depreciation cost was twice that of the material costs.at 10% of the total costs. Similar to Geometry𝐺𝐺2𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑚𝑚𝑚𝑚𝑀𝑀ℎ𝑖𝑖𝑎𝑎𝑠𝑠 𝐺𝐺1𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑏𝑏𝑠𝑠𝑠𝑠𝑜𝑜𝑠𝑠𝑀𝑀𝑖𝑖𝑚𝑚𝑠𝑠𝑖𝑖𝑜𝑜𝑎𝑎𝐶𝐶𝑠𝑠𝑜𝑜𝑠𝑠𝑠𝑠𝑠𝑠Figure 13: Total Costs vs. Build Cycles for PowderFigure 14: Cost for Workflow Activities forReuse for GeometryGeometrywith Virgin Powder𝐺𝐺25.3. Summary of Results𝐺𝐺2The costing for these two example parts revealed that the machine cost,was the most pervasive cost for both builds. This is due to the volume of the geometries and 𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑚𝑚𝑚𝑚𝑀𝑀ℎ𝑖𝑖𝑎𝑎𝑠𝑠 support structures being consolidated during processing, their corresponding build heights, and the build speeds at which the DMLS machine can melt the given material.

The labor activities showed no significant cost fluctuation pertaining toand𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏𝑏𝑏𝑜𝑜𝑏𝑏, 𝐶𝐶𝑠𝑠𝑠𝑠𝑠𝑠𝑢𝑢𝑠𝑠, 𝐶𝐶𝑜𝑜𝑠𝑠𝑚𝑚𝑜𝑜𝑅𝑅𝑚𝑚𝑏𝑏,𝐶𝐶𝑠𝑠𝑢𝑢𝑏𝑏𝑠𝑠𝑠𝑠𝑜𝑜𝑚𝑚𝑠𝑠𝑠𝑠2023,among the example parts. This is because these activities are standardized procedures with average completion times based on the skill of the AM operator and independent of the geometries in the build. While can vary for complex geometries requiring support structures, these costs ranged between 3-11% of the costs in all of the builds, due to 𝐶𝐶𝑠𝑠𝑜𝑜𝑠𝑠𝑠𝑠𝑠𝑠 larger costs being attributed to machine time, material, and deprecation.and𝐶𝐶𝑠𝑠𝑜𝑜𝑠𝑠𝑠𝑠Depreciation was one of the largest costs for Geometriesand.

It was at most 42% and its use of Ti64 requiring nearlybut only 11% in of the total costs for powder, the relatively larger depreciation cost in comparison to twice as much powder to fill the feed bed due to differences in tap densities and part build heights, along with the Ti64 powder being nearly seven times more expensive than GP1. The machine cost had a relatively higher percentage in GP1 because of its 20 micron layer thickness build speed, which was half the speed when using Ti64. In both examples, when and 7.2 using a powder with a of 10 build cycles, the depreciation cost was greater than twice the 3 h/cm cost of the melted material. One interpretation for this result is that the utilized build envelope (i.e., volume packing) [53] is uneconomical for the given build since the surrounding un-melted powder is being put at risk of contamination and degradation, and thus more costly than the built-up parts produced in the AM process.

Given that the depreciation cost is calculated as the difference between the mass of the un-melted powder and the mass of powder melted for the parts, a highly packed build tray with a high quantity and large mass of produced parts could reduce the overall depreciation costs since more of the surrounding powder would be melted and consumed during the build job. Less un-melted powder would remain after the build and thus, the costs for the material could be more economical than the depreciation, however post- processing costs may increase as a result of the additional part quantities.. When looking specifically at 𝐺𝐺2𝐺𝐺1is due to 𝐺𝐺1𝐺𝐺1𝐺𝐺2𝐺𝐺2𝐺𝐺1𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚5.4. Model LimitationsUncertainties in the model are the exact number of reuses permitted for each materialalloy, an open and active area of research in the metal AM community.

Model LimitationsUncertainties in the model are the exact number of reuses permitted for each materialalloy, an open and active area of research in the metal AM community. The parametric analysis accounted for 4 different maximum build cycle limits, but literature and standards are limited that provide recommendations on the exact extent for which a reused metal powder can be used to produce AM parts in functional engineering applications (e.g., aerospace, biomedical, etc.). This model did not consider the use of alternative powder feedstocks not officially offered by the OEM, which may have lowered costs and altered process parameters. However, the use of unauthorized materials run the risk of voiding AM machine warranty agreements. Since the powder is selectively melted with each build job, the remaining un-melted powder lot would require additional powder to be added in order to reproduce the same build job.

Since the powder is selectively melted with each build job, the remaining un-melted powder lot would require additional powder to be added in order to reproduce the same build job. Thus, a cost would have to be allocated for adding virgin powder to a reused powder lot in production. Post- processing accounted for the removal of support structures; however, functional components may call for more sophisticated processes such as annealing, shot peening, or CT (computed tomography) scanning in order to validate the integrity of the part. While these examples considered DMLS, costing for EBM will have differences due to variations in PBF technology and required labor activities. In addition, the empirical waste factors in calculating the part mass were specific to an EOSINT M280 and may not be applicable to other PBF processes. Finally, costs for the re-design and engineering of a pre-existing component in order to be made using AM were not captured in this model.20246.

Finally, costs for the re-design and engineering of a pre-existing component in order to be made using AM were not captured in this model.20246. Closing RemarksA cost model was created for reused metal powder feedstocks using SOYD depreciationto account for costs pertaining to the diminishing and depletion of a metal powder feedstock’s properties in PBF. This depreciation model was implemented in a generic activity-based costing model for PBF. A case study was presented using Ti64 and GP1 parts produced using DMLS. Costs for individual parts showed that were the top two costs, whereas labor activities were relatively stagnate and less costly. With reference to cost 𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑚𝑚𝑚𝑚𝑀𝑀ℎ𝑖𝑖𝑎𝑎𝑠𝑠 models with fixed material costs, the depreciation costs due to lost value from reusing a metal powder feedstock added an additional 13-75% in cost for build jobs with virgin powders using Ti64, or 3-11% for GP1.

Upon exceeding 13 build cycles in Ti64, 11 for GP1, the total costs for build jobs achieved a cost savings of 38% or 5% when using a highly reused powder or powder feedstock exceeding its useful life ( feedstock will vary based on material type, PBF technology, and the functional application of the resulting component.and𝐶𝐶𝑏𝑏𝑢𝑢𝑖𝑖𝑏𝑏𝑏𝑏−𝑏𝑏𝑠𝑠𝑠𝑠𝑜𝑜𝑠𝑠𝑀𝑀𝑖𝑖𝑚𝑚𝑠𝑠𝑖𝑖𝑜𝑜𝑎𝑎). The practicality of using a highly reused powder𝑈𝑈𝑚𝑚𝑚𝑚𝑚𝑚Future improvements to this model could explore more detailed costing options for post- processing. The financial implications of volume packing and adding virgin powder to a reused powder lot could be explored to better understand building scenarios that are most applicable to high volume production in industry. Although this study considered the cost for two individually produced components, the supply chain and economics of mass production could be modeled to study the impact of accounting for reused powders.

One improvement would be the creation of a standardized metric for measuring the overall quality and reusability of the powder feedstock in a PBF technology to aid in formalizing the financial value and acceptability for a reused powder. Finally, further material science research on metal powder feedstocks for AM can help broaden the literature that quantifies mechanical properties and the engineering limitations of these materials as they are continually adopted in full-functional components.AcknowledgmentsThe authors would like to thank Richard Martukanitz, Todd Palmer, Kenneth Meinert, Abdalla Nassar, and all the CIMP-3D staff for their assistance and support in this study. We would also like to recognize Eduard Romaguera for his early efforts on researching AM cost literature and deriving some of the initial equations for this model.References[1] ASTM F2792 - 12a Standard Terminology for Additive Manufacturing Technologies.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceCotton Fibers in 3D PrintingAlexa Kearns*, Nasim Farahbakhsh*, Richard Venditti†, and Jesse Jur*Department of Textile Engineering, Chemistry, and Science †Department of Forest Biomaterials North Carolina State University, Raleigh, NC 27606AbstractThis work explores the materials challenges of cotton-loaded with polymer composites toward sustainable solutions in 3D printed free forms. A key step toward composite filament development is the reduction in size of the original cotton fibers. Mechanical processing of the cotton is introduced as a means of reducing the size of cotton fibers to form a material of an ultra- high aspect ratio (>250) structure that is nanometers in diameter and micrometers in length.

Mechanical advantages are low density polyethylene loaded with the high aspect ratio cotton and is observed to maintain a mechanically robust material at loading up to 40 wt%. In addition, attempts to print with 25 wt% cotton fillers (~10-15 aspect ratio) in LDPE is demonstrated. Finally, considerations to processing challenges from a sustainable and practical viewpoint are provided.Keywords 3D printing, Cotton, Pulverized cotton, low-density polyethylene, nanofibrillated cellulose (NFC)IntroductionFreeform fabrication and rapid prototyping by additive manufacturing (AM) is an emerging field for creating products and parts for specific applications that have the potential to reduce manufacturing costs. The importance of reducing the consumption of energy becomes evident when as much as one third of U.S. energy cost are devoted to the manufacturing sector [1].

The importance of reducing the consumption of energy becomes evident when as much as one third of U.S. energy cost are devoted to the manufacturing sector [1]. Reducing the amount of energy consumed lessens the reliance on fossil based fuels and other forms of energy harvesting that can negatively impact the environment. However, because of the time necessary to create a product, the energy consumption between AM and traditional manufacturing remains relatively equivalent [2].Adding different plant-based or cellulose-based materials into traditional filaments can create unique properties that add to the sustainability of AM. Materials such as algae, bamboo, coffee grounds, hemp, and woodfill are commercially available. The mechanical properties tend to be weaker than traditional filaments, such as acrylonitrile butadiene styrene (ABS) and Poly(lactic acid) (PLA).

The mechanical properties tend to be weaker than traditional filaments, such as acrylonitrile butadiene styrene (ABS) and Poly(lactic acid) (PLA). However, when attempting to mimic natural elements (like a pine cone) the properties of natural based filaments exceeded that of synthetic filaments, especially with moisture uptake and curvature detail [3]. These natural filaments tend to contain up to 40% fillers that originated as pulp or cellulose-based nanofibers. These nanoparticles are able to seamlessly blend into powdered synthetic filaments because of a lager surface area, compared the volume of the particles [4].Cotton based filaments provide a unique set of properties that makes it of interest in 3D printing. For example, cotton is a good conductor of heat, which would allow for a uniform heating of the filament during printing. Additionally, unlike wood, cotton gains strength when wet and therefore can be used in more applications where moisture is involved [5].

Additionally, unlike wood, cotton gains strength when wet and therefore can be used in more applications where moisture is involved [5]. The purity and899molecular weight are both higher in cotton than in wood products and when processed the crystallinity of cellulose does not decrease [6]. Finally, the size of cotton can be scaled to larger extents in as compared to wood pulp, thus allowing for the high aspect ratio of sized cotton to achieve a composite-like enhancements in the printed freeform. The aim of this work is to first demonstrate procedure for the sizing of cotton and effects of the cotton when used as a plant-based filler in 3D print polymers.Experimental DesignThe cotton materials used in these experiments detailed in this research was obtained as pulverized cotton (pCot) from Solvaira Specialties. This cotton is derived from recycled t-shirts and ground into micro-sized fibers, 350 µm in length and 20-40 µm in width.

This cotton is derived from recycled t-shirts and ground into micro-sized fibers, 350 µm in length and 20-40 µm in width. In comparison, cotton fiber as a typical dimension of X mm in length and 20 µm in diameter. A microgrinding procedure is used to further alter the size of the pCot to a larger aspect ratio. This is conducted using a Masuko Grinder microfluidizer in which a 3 wt % solution of pCot (in water) passes between two rotating stone plates at a gap of -.28 mm rotating at 1500 rpm [7]. The friction between the stones causes the solution to grind the cotton into nanofibrillated cotton (NFC) particles. As previously demonstrated, nine passes of the solution in the microfluidizer results in a uniformly reduced size of cotton. The NFC is then freeze dried in a two-step process, first at -80°C in a glass jar and then at -40°C in a Labconco freeze dryer.

The NFC is then freeze dried in a two-step process, first at -80°C in a glass jar and then at -40°C in a Labconco freeze dryer. Once the fully dried, the NFC is then processed in a blender and shaken through a sieve to produce a powdered product.At this point two different avenues are pursued low density polyethylene (LDPE) composites with the originated pCot material and also with the NFC material. The LDPE used in these experiments were obtained from Dow Chemical (AsSpun 6850A) and used as received. To compound the polymer with the pCot and NFC a DSM, Xplore, Micro 15 cc Twin Screw Compounder is used at heating of 170 °C, screw speed of 50 rpm, and a mixing time of 10 min. Compounded samples of up to 20 wt. % cotton material were produced.

Compounded samples of up to 20 wt. % cotton material were produced. In addition, commercial- grade composite materials were fabricated and produced into pellet sizes with 25% pCOT and LDPE by Techmer PM.Using a Noztek pro HT filament extruder with a 0.05mm nozzle size the LDPE- pCot pellets were extruded at 185°C then set to rest until cooled. The tensile strength of the filament was tested to compare against LDPE filament without the addition of pCot. Following ASTM D368-14, five different five inch segments of the filament were pulled in tension at 50.82 mm/min. The remaining cooled filament was loaded into a Lulzbot TAZ5 3D printer and extruded at 190°C creating a 127 x 5.1 x 35.6 mm dog bone shape. The print’s mechanical properties are then tested to compare to other natural filaments and LDPE without pCot.900Results and DiscussionFilament Extrusion.

The print’s mechanical properties are then tested to compare to other natural filaments and LDPE without pCot.900Results and DiscussionFilament Extrusion. This effort explores the differences between the LDPE extruded from the Noztek pro HT with and without cotton and define the best extrusion conditions. In these experiments, an initial amount of filament was extruded for approximately 1 min prior to the collection of the samples for collection. Microscope images of the LDPE extruded at 145°C, 165°C, 185°C, and 200°C are provided in Figure 1a, showing the average diameter decreasing as the temperature increased (Table 1). The mechanical strength of the virgin LDPE compared to LDPE with pCot. The strength of the Virgin LDPE is substantially higher than the strength of LDPE with pCot. This is due to the high loading of the cotton fibers decreasing the strength.

The strength of the Virgin LDPE is substantially higher than the strength of LDPE with pCot. This is due to the high loading of the cotton fibers decreasing the strength. Additionally, moisture could have infiltrated the composite during the extrusion process.110 ∘C120 ∘C 130 ∘C4.57 mm4.25 mm3.86 mmFigure 1. Comparison of LDPE filament extruded at temperatures ranging from 100-130 °C in a Noztek pro HT.Table 1. Diameters of filaments extruded with the Noztek pro HT.Virgin LDPELDPE with pCotTemperature (°C) 145 165 185 200Diameter (mm) 4.58 4.34 3.3 3.3Temperature (°C) 110 120 130Diameter (mm) 4.57 5.25 3.86901As a comparison, the LDPE extruded with 25 wt% pCOT at the same temperatures are provided in Figure 2. Again, the average diameter is observed to decrease with an increase in temperature and the results are provided in Table 1.

Again, the average diameter is observed to decrease with an increase in temperature and the results are provided in Table 1. In comparison to the virgin LDPE filament, the composite filament is demonstrated to have a rougher surface and similar diameter. The similar diameter is due to the nozzle head on the Noztek pro being the same size. The virgin LDPE was slightly larger due to having no particles impeding the flow.4.584.343.303.30Figure 2. Comparison of the LDPE filament with 25 wt% pulverized cotton (pCOT) extruded at a temperature range of 145-200 °C in a Noztek pro HT.Mechanical analysis of filaments was performed, with the stress-strain curves presented in Figure 3 and key mechanical values provided in Table 1. With respect to a normalized diameter, the filament extruded at 145 °C had the highest strength value, but proved to be inconsistent with regards to the extruded diameter and quality.

With respect to a normalized diameter, the filament extruded at 145 °C had the highest strength value, but proved to be inconsistent with regards to the extruded diameter and quality. Failing at about 45 N, 185 1 was determined to be strongest and most consistent filament. However, all of the extruded filaments with an LDPE pCot mixture were weaker and could withstand less strain than the virgin LDPE filaments.9023D printing of LDPE/pCOT. 3D printing was attempted using the LDPE/pCOT filament extruded at 185 °C. The typical print temperatures of LDPE is most often performed at a range of5090458070)40)NN60( e c r o F353025( e c r o F50403020200.51.01.52.02.53.03.54.04.50246810Strain (mm)Strain (mm)Figure 3. Strain curves of the four tested temperatures for LDPE pCot filament extrusion and b. LDPE pellets without pCot extruded at differing temperatures.160-180 °C.

Strain curves of the four tested temperatures for LDPE pCot filament extrusion and b. LDPE pellets without pCot extruded at differing temperatures.160-180 °C. Ideally, a broader range of extrusion temperatures would be selected for printing in this work, due to observations in the change in extrusion temperature observed with the inclusion of pCOT in the LDPE. Still, 3D printing is limited to a high temperature of 200 °C due to thermal limitations to the cotton. The best success of the 3D printing was observed at a temperature ofabFigure 4. Image of a. the printing process of LDPE pCot and b. the initial results of the print. Developing NFC190C (Figure 4). In the printing shown, the print surface was unheated and a glue stick was used to assist in the bonding. The result of the printing showed a successful initial layer of printing, but a multiple layered-structure was unable to be fabricated.

The result of the printing showed a successful initial layer of printing, but a multiple layered-structure was unable to be fabricated. Printed virgin LDPE and LDPE with pCot was extruded from the 3D printer and the mechanical strengths were tested. Figure 5 shows LDPE to be the stronger print, which is due to not having a filler polymer. Additionally, the LDPE pCot pellets displayed signs of containing moisture which can detract from the strength.903363432LDPE LDPE + pCot)N30(e c r o F2826242220024681012Strain (mm)Figure 5. Strain curve of the virgin LDPE compared to the strain curve of LDPE with pCot.Sizing of cotton to higher aspect ratios. This purpose of this effort is to overview our team’s prior effort toward further sizing of cotton to larger aspect ratios.[7] As outlined in the experimental section, the route taken to achieve a larger aspect ratio was through microgrinding.

Our results have shown that the nine passes of the pCot through the microgrinding plates, results in a uniform nanofibrilatted cotton structure, has aspect ratios of >250. An SEM and TEM image analysis is of the pCot and NFC materials is provided in Figure 6.Figure 6. (left) SEM image of pCot before microgrinding. (middle) SEM image of the Cot after 9 microgringinComparison of the LDPE filament with 25 wt% pulverized cotton (pCOT) extruded at a temperature range of 145-200 °C in a Noztek pro HT.ConclusionThis works examines the creation of nano-sized cotton particles and the integration into additive manufacturing. SEM and TEM images were taken to illustrate the structure of the cotton904after two, four, and nine passes in a microgrinder. These images show greater fibrillation of the cotton with more passes and developed structures ranging 5-15 nm in length.With the help of Techmer PM LDPE and pCot were compounded to create a pellet.

An extrusion at 185°C the filament created was strong enough to be printed in a Lulzbot TAZ5 3D printer with success. The results produced a weak print that had trouble binding with other printed layers, however manipulating the heating and print bed should be a remedy. Compounding PLA and NFC proved to be a challenge with a 3:1 weight percent ratio. This work shows that printing with a cotton cellulose material is possible and has the potential to be a mass produced product.AcknowledgementsThe authors would like to acknowledge Cotton Inc. for their financial and visionary support. The help of WestRock for the processing of the NFC was immensely helpful. For the supply of pCot Solvaira Specialty help has also be appreciated. Additionally, we would like the thank Wissam Farhat (N.C. State) for the assistance in the compounding and freeze drying processes.References [1] M. Mani, K. W. Lyons, and S. K. Gupta, “Sustainability characterization for additive[2]manufacturing,” J.

EATING COMPLEX HOLLOW METAL GEOMETRIES USING ADDITIVE MANUFACTURING AND ELECTROFORMINGD. L. McCarthy and C. B. WilliamsDREAMS Lab, Department of Mechanical Engineering Virginia Tech, Blacksburg, VA 24061ABSTRACTAdditive manufacturing introduces a new design paradigm that allows the fabrication of geometrically complex parts that cannot be produced by traditional manufacturing and assembly methods. In this paper, the authors investigate the combination of laser sintering with an electroforming process using electroless nickel plating to produce complex, thin-walled, hollow, metal geometries. The resulting geometries cannot be produced directly with other additive manufacturing systems. The resulting process is used to produce a cellular nickel structure featuring 800µm walls that is 65 vol% air from a polyamide substrate with 3mm pores.Keywords: Additive Manufacturing, Electroless Plating, Electroforming, Laser Sintering1. INTRODUCTION1.1.

INTRODUCTION1.1. MotivationPast work has shown that the use of hollow members in applications like high strength-to- weight ratio sandwich panel structures can have significant benefits, including improvements to the peak compressive collapse strength and inelastic buckling resistance [1]. In general, sandwiched trussed structures are typically made with a folding and brazing operation on a diamond perforated metal sheet. In structures featuring hollow tubes, the same approach is taken, but instead by bending an arrayed series of hollow metal tubes. In both cases, the resulting structures are limited to planar macrostructures, and a limited set of mesostructure geometries that can be shaped via bending.To address these geometric constraints, the authors look to additive manufacturing (AM) as a means of providing the freedom necessary for a designer to create complex hollow metal cellular artifacts with a designed mesostructure.

Combining the benefits of hollow structures with the design complexity provided by additive manufacturing increases the range of potential applications, including more advanced high strength-to-weight ratio structures. In addition, as hollow members enable the transmission of fluids, the resultant high-surface area structures could be used as heat exchangers.1.2. Metal Additive ManufacturingAs the mechanical properties and heat conductivity of metals make them ideal for these applications, metal AM processes are candidates for providing the desired complex hollow geometries. Direct metal AM processes (e.g., direct metal laser sintering, selective laser melting, and electron beam melting) tend to provide a rough surface, which may cause problems for applications involving fluid flow.

Additionally, the majority of the metals that can be used in these processes have poor heat conductivity [2, 3], excepting aluminum (which provides its own108difficulties due to poor flowability and high reflectivity [4]). High heat conductivity metals like copper could theoretically be used, but that conductivity actually acts against the system, lowering part accuracy and increasing part growth due to the uncontrollable melt pool and solidification process [5]. Both copper and aluminum also experience major oxidation issues during sintering, which restricts diffusion and reduces the wettability of the surface [4].Indirect metal processes (e.g., two-phase selective laser sintering and three-dimensional printing) require involved post-processing that can be difficult for complex structures. Small- scale truss structures, for instance, are not producible with indirect metal 3D printing because of the weakness of the green part after printing before post-processing.

Small- scale truss structures, for instance, are not producible with indirect metal 3D printing because of the weakness of the green part after printing before post-processing. Attempting to produce a truss structure with hollow members is impossible primarily due to powder removal requirements while the part is still weak. Additionally, infiltrating complex thin walls with another metal during post-processing is difficult.Both direct and indirect powder bed fusion processes like those described require powder removal which is difficult or impossible for long, hollow channels in complex structures. Additionally, the available resolutions of these processes do not allow the creation of thin walls (on the order of 100µm) with reliable structural integrity.1.3. Hybrid Manufacturing ApproachWhile current metal AM systems might be able to produce complex hollow metal geometries, the resultant geometries are limited by the laser spot size.

Hybrid Manufacturing ApproachWhile current metal AM systems might be able to produce complex hollow metal geometries, the resultant geometries are limited by the laser spot size. For example, the wall thickness of hollow geometries cannot be smaller than the laser spot diameter. To provide more flexibility in the design and realization of complex hollow geometries, the authors propose a hybrid electroforming approach wherein additively manufactured thermoplastics are used as sacrificial substrates that are removed once they have been plated in metal, leaving behind the thin walls of the metal plate. A generalized representation of the proposed hybrid process is shown in Figure 1.Figure 1. Diagram of the generalized proposed process by sub-function for producing complex, hollow, metal geometries.Several groups have used similar approaches in the past (though most only in part).

Diagram of the generalized proposed process by sub-function for producing complex, hollow, metal geometries.Several groups have used similar approaches in the past (though most only in part). Chandrasekhar [6], Saleh [7], Zhou [8], and their respective coauthors all used an approach that used stereolithography (SL) or laser sintering (LS), electroless plating, and electroplating to produce plated parts from AM substrates. Their goals were to improve the mechanical properties of the AM parts by adding a layer of metal coat. These studies addressed only simple geometries (i.e., airfoils and tensile testing “dog bones”) without the complex internal features desired for this work. Additionally, they do not consider substrate removal procedures.109Markkula and coauthors [9] investigated the use of fused deposition modeling (FDM), electroless plating, and electroplating to produce high strength-to-weight ratio sandwich trusses.

They tested pyramidal, tetrahedral, and strut-reinforced tetrahedral (SRT) truss formations and compared their compressive properties. Despite being truss structures, the process was simplified by producing only a single truss cell for each test, which does not require consideration of geometric dependencies in the plating process. This procedure also did not include any substrate removal process.Monzón and coauthors [10] used FDM, electroless plating, electroplating, and mechanical separation to produce electrically conductive patterns for various applications. They have also explored using FDM in an electroforming context to make low quality injection molds for rapid tooling [11]. Their research focused on taking advantage of the internal stresses in FDM parts to reduce the adhesion between the electroless coat and AM substrate.

Their research focused on taking advantage of the internal stresses in FDM parts to reduce the adhesion between the electroless coat and AM substrate. The parts were simple, allowing for mechanical separation of the AM substrate from the metal coat while preserving both the substrate and the electroform.The rapid tooling of electric discharge machining (EDM) electrodes accounts for the majority of research involving the use of AM substrates in a complete electroforming process. Most researchers select SL as the AM process of choice because of its high resolution and smooth surface finish. Arthur and coauthors [12] used conductive paint, electroplating, and tested both thermal (burnout) and chemical (etching) substrate removal. Yang and Leu [13, 14] used electroless plating, electroplating, and thermal substrate removal.

Yang and Leu [13, 14] used electroless plating, electroplating, and thermal substrate removal. These groups concluded that thermal substrate removal has significant negative effects on the quality of the electroform due to the volatility of the SL photopolymers. Work by Gillot [15], Dimla [16], and their respective coauthors concluded that electroplating is too geometrically dependent to be used in this process, showing considerable non-uniformity, affecting the conductivity of the electrode.The design of a new AM-inspired process for creating microlattice structures from SL photopolymers [17, 18] led to a process that combined an SL-like process with electroless plating and chemical etching from Schaedler and coauthors [19] to produce ultralight metallic microlattices. The resultant geometries were limited to straight-lined truss structures with limited dimensions, not taking advantage of the design freedom provided by AM.

The resultant geometries were limited to straight-lined truss structures with limited dimensions, not taking advantage of the design freedom provided by AM. Additionally, since the properties of these structures are specific to the nanocrystalline electrodeposition of the nickel used [20], producing copper structures to improve heat and electrical conductivity requires a different process that would likely not be able to replicate the structural recovery properties of the nanocrystalline nickel.1.4. Design GoalsWhile past hybrid AM/metal plating processes have successfully produced high quality metal coatings on printed polymers, considerable effort remains to finalize the proposed hybrid process chain (Figure 1). A successful hybrid process will be able to produce complex, hollow, metal geometries with the full design freedom provided by AM.

A successful hybrid process will be able to produce complex, hollow, metal geometries with the full design freedom provided by AM. The manufactured structures must be exclusively metal (preferably with high heat conductivity), be self-supporting, and be producible with any level of design complexity. Ideally, the developed process would be relatively inexpensive and be possible in a lab setting, without requiring outsourcing or extensive knowledge of complicated processes.110Section 2 discusses process selection, using a morphological chart to investigate process combinations and eliminate unhelpful solutions for each process sub-function. Section 3 details the procedure used to produce a complex, hollow, metal geometry. Section 4 provides the results and analysis of a part after completing the process. Section 5 concludes the work with the contributions and limitations as well as the recommendations for future work.2.

Section 5 concludes the work with the contributions and limitations as well as the recommendations for future work.2. PROCESS SELECTIONFigure 2 shows a morphological chart based on the sub-functions presented in the generalized process diagram in Figure 1. It is populated with potential solutions for each sub- function. Identifying combinations with one solution from each sub-function helps to provide process alternatives that might otherwise be missed.

Identifying combinations with one solution from each sub-function helps to provide process alternatives that might otherwise be missed. In this section, these potential solutions are investigated and most are eliminated to narrow the experimental scope.Additive ManufacturingConductivizingMetal CoatingSubstrateRemovalSolutionsSub-FunctionsSelective LaserSintering (SLS)Fused Deposition Modeling (FDM)Direct Inkjet 3DPrinting (3DP)Stereolithography (SL)AgNO3 to Ag2S ConversionConductive Paint(Silvaspray™)ElectrolessPlating Burnout(Thermal)Etching(Chemical)Separation(Mechanical)NoneElectroplatingVacuum MetallizationMetalSprayingCathode SputteringElectroless PlatingFigure 2. Morphological chart for a process capable of creating complex, hollow, metal geometries.2.1. Substrate RemovalSubstrate removal can be done through thermal means (e.g., burning out a thermoplastic), chemical means (e.g., etching away a photopolymer), or mechanical means (e.g., forcibly separating the substrate).

Mechanical separation can only be done for very basic geometries and is not possible with the complex geometries investigated in this work, thus eliminating it from consideration. A thermal burnout process is ideal for materials that melt, which is why it is used in the investment casting process. It therefore works particularly well for waxes and thermoplastics (like ABS plastic and nylon, used in the FDM and SLS processes, respectively). This does not hold true for photopolymers (used in direct 3DP and stereolithography), which are thermosetting plastics that do not melt, instead violently expanding during burnout. While some research suggests that narrow channels in complex parts made from photopolymers can be removed after plating via chemical etching [19], the required etch changes with each photopolymer, of which there are many. The Objet photopolymers used in the direct 3DP process are not well-characterized, thus eliminating chemically etching as an option for 3DP parts.

The Objet photopolymers used in the direct 3DP process are not well-characterized, thus eliminating chemically etching as an option for 3DP parts. While etching is a possibility for SL parts, complexity is an issue. Etching chemicals must have access to the SL part. With a complex part that may include only one or two small entrances for etching chemicals, the etching process could take several days or weeks, assuming it is possible at all. Thermal substrate removal was therefore selected as the most reasonable substrate removal solution.1112.2. Additive ManufacturingSL and direct 3DP parts are made of photopolymers. While the parts have very high resolution, they require extensive manual support removal after fabrication, which can be difficult or impossible to remove when producing complex structures. Additionally, testing showed that photopolymers violently expand as they are heated to the point of degradation.

Additionally, testing showed that photopolymers violently expand as they are heated to the point of degradation. This expansion causes a great deal of internal pressure in the electroform, which catastrophically ruptured copper plating that was over 750µm thick, as shown in Figure 3. This chaotic thermal expansion eliminated both SL and 3DP from consideration in this procedure.(b) Figure 3. Burnout test of basic 3DP photopolymer part (with 20mm dimensions) after (a) copper electroplating, (b) burnout, (c) cleaning showing plate rupture.(a)(c)FDM produces parts from thermoplastics, typically ABS. ABS is the most commonly electroplated and electrolessly plated type of polymer in industry because of the simplicity of etching before electroless nickel plating (ENP) [21]. As a thermoplastic, ABS provides the benefit of being able to melt during a potential burnout process.

As a thermoplastic, ABS provides the benefit of being able to melt during a potential burnout process. Unfortunately, the FDM process provides clear delineations between layers due to relatively large layer thicknesses (on the order of 254µm), showing a “stair-stepping” effect on curves and angles. This effect precludes the use of the FDM process in the creation of complex geometries because angles and curves are a requirement for design freedom. As such, use of the FDM process was removed from consideration.SLS most commonly produces parts from DuraForm™ PA (polyamide) powder, also known as Nylon 12. As a thermoplastic, polyamide is also a good candidate for a polymer burnout processes. Unfortunately, polyamide is rarely used for plating processes because it is hygroscopic, leading to potential warping when absorbing moisture during plating, so plating on nylon is not well-researched.

As such, testing was performed to ensure that SLS parts could successfully be plated and burned out. Preliminary tests showed that SLS parts accept a conductive layer and a copper electroplate with a wax coating. The preliminary wax infiltration mitigated hygroscopy issues by pre-stressing the part, removing the possibility for detrimental structural changes during plating. The parts were successfully burned out from a 500µm thick copper electroplate without rupturing the plate. SLS was therefore selected as the most reasonable solution for the AM sub-function.2.3. Metal Coating and ConductivizingVacuum Metallization, Metal Spraying, and Cathode SputteringSeveral technologies exist for coating metals onto polymer parts. Three widely used technologies are vacuum metallization, metal spraying, and cathode sputtering. In vacuum metallization, evaporated metal is condensed on the substrate.

Three widely used technologies are vacuum metallization, metal spraying, and cathode sputtering. In vacuum metallization, evaporated metal is condensed on the substrate. In metal spraying, metal is112atomized and propelled toward the substrate by a high velocity gas jet. In cathode sputtering, a high electrical potential disintegrate the surface of the source metal, projecting the disintegrated atoms onto the substrate. These three technologies each have high capital costs and can be very expensive for small batches of parts. Most importantly, they are each geometrically dependent, limiting their applicability for complex geometries. As such, they were eliminated from consideration in this process.Electroplating and ConductivizingElectroplating, also known as electrolytic plating, is a process in which an electrical current moves metal ions through an acid bath from a source metal (the anode) to the substrate (the cathode).

Electroplating requires that the substrate have a conductive surface to complete the circuit. Conductivizing is typically done with either a base coat of electroless plating (usually nickel) or conductive silver paint. Preliminary tests showed that conductive paint provided a consistent conductive layer at a lower cost than ENP.Electroplating is also a geometrically dependent process (the plate is thicker the closer the feature is to the anode), however the low cost and ease of use made it worthwhile to investigate. Extensive testing with copper electroplating showed that the electroplating process is considerably too geometrically dependent to be used in this process, as shown in Figure 4.Figure 4.

Extensive testing with copper electroplating showed that the electroplating process is considerably too geometrically dependent to be used in this process, as shown in Figure 4.Figure 4. Cross-section of an electroplated part with a close up of the lower left strut showing the poor plate quality and large gradient from exterior (facing the anode) to interior (facing away from the anode).The parts showed a considerable gradient from the exterior (closer to the anodes) to the interior. Even at high currents (which increase penetration into the part) with large entry pores, the copper did not plate all the way to the interior of the 48mm x 48mm x 48mm part, leaving an unplated region on the interior features of the central struts. Additionally, the plating quality was so low and unpredictable that the resultant parts would be useless for most applications.

Additionally, the plating quality was so low and unpredictable that the resultant parts would be useless for most applications. Due to these issues, electroplating was eliminated as a potential solution for the metal coating sub- function.Electroless PlatingElectroless plating, also known as autocatalytic plating, is powered through a chemical reaction, as opposed to an electrical circuit. No anode is required as a source metal because the source metal is already present in the plating solution. This also means that the source metal in the solution must be periodically replenished as it is used. Deposition is provided by metal ions in the solution that are reduced to their metallic state in the presence of a reducing agent. ENP is most common and best understood. While electroless processes are more expensive than electroplating processes, they provide two advantages that are integral to this work.

ENP is most common and best understood. While electroless processes are more expensive than electroplating processes, they provide two advantages that are integral to this work. First, they113can plate on non-conductive surfaces with the proper pretreatment. Second, and most important, they do not depend on the geometry of the workpiece and provide an even plating coat regardless of the complexity of the part. Because of these advantages, electroless plating was selected as the solution for the metal coating sub-function.3. PROCEDUREFrom the review presented in Section 2, the selected morphological combination is shown in Figure 5. The proposed procedure begins with fabrication of a DuraForm™ polyamide part using SLS (Sinterstation 2500plus), followed by cleaning of excess powder with compressed air, then high pressure water cleaning. The part is then dried and degreased with acetone.

The part is then dried and degreased with acetone. The part is then etched with a sodium hydroxide (NaOH) solution [22], then undergoes a room temperature metallization (RTM) procedure for electrolessly plating a base coat of nickel onto the polymer (RTM kit for plastics from Transene, Inc.). After rinsing, the part is then submerged in a larger electroless nickel kit designed to plate nickel onto a base coat of nickel (ENP kit from Caswell, Inc.).Figure 5. Updated morphological chart for process generation indicating which solutions are not feasible and indicating the selected morphological combination.Because the plating nickel is present in the solution (as opposed to being in an anode), periodic chemical replenishment is necessary. Since the ability of electroless plating to produce thick coats is not well-researched, this experiment was also used to investigate if very thick coats were possible.

Since the ability of electroless plating to produce thick coats is not well-researched, this experiment was also used to investigate if very thick coats were possible. As such, the part was allowed to plate until the bath was “killed”, having gradually degraded during the reaction. The resultant plate thickness was about 800µm, showing that thick coats are possible with ENP. Such a thick coat required chemical replenishments every 100 minutes.The plated part underwent a flash burnout procedure for substrate removal. In this process, the part is heated well above the temperature at which most of the mass is degraded in a very short period, essentially shocking the material into degradation before it is allowed to expand significantly. A custom-built gas fired kiln in Virginia Tech’s Kroehling Advanced Materials Foundry was used for this procedure.

A custom-built gas fired kiln in Virginia Tech’s Kroehling Advanced Materials Foundry was used for this procedure. The part was placed in the furnace, which ramped from room temperature up to 750°C (180°C greater than the temperature at which 96% of the polyamide should degrade [23]) in about five minutes. The part was left at this temperature for 15 minutes, then allowed to cool for about 20 minutes before opening the furnace. Additional procedural details are presented in [24].1144. RESULTS AND ANALYSISTo evaluate the ENP process, tests were conducted on a 3mm pore diameter part. The full size (48mm cubed) part underwent the fabrication, cleaning, etching, and RTM base coat. The RTM ENP solution produced large bubbles (on the order of about 3mm in diameter) that had a tendency to become trapped in the interior of the part, requiring agitation to continue plating. This process deposited a nickel plate between 5 and 10µm thick.

This process deposited a nickel plate between 5 and 10µm thick. The part was then bisected, which confirmed that there was consistent plating throughout. The part was then further sectioned down to dimensions of 21mm x 22.5mm x 25.5mm.The progress of the main plate over time was evaluated by analyzing images taken at specific times, providing a coarse estimate of the plating rate in the ENP bath. Examples of these images are shown in Figure 6.(a) (b) (c)(d)Figure 6. 3mm pore diameter part section during main ENP coating at (a) 0h50m, (b) 6h50m, (c) 11h50m, and (d) 15h10m.After 24 hours of constant plating, the part had developed a plate thickness of about 800µm. The part plated at an average rate of 39µm per hour (via linear approximation; R2 = 0.9415).

The part plated at an average rate of 39µm per hour (via linear approximation; R2 = 0.9415). While the rate of plating is relatively linear over short ranges, for thick coatings, the surface area changes significantly enough that the rate of plate growth may decrease slightly (explaining the discrepancy between the plate thickness after 24 hours and the plating rate during the first 15 hours of plating). This plating rate is dependent upon the plating duration, the surface area of the part, and the volume of the ENP bath. The chemical proportions of the bath are determined by the surface area of the part, so while the plating rate would normally be constant no matter the surface area of the part, a significant surface area change during plating will slow the reaction. The degradation of the bath over time can have a similar effect.

The degradation of the bath over time can have a similar effect. Larger volume baths will allow the part to plate faster.After plating, the part was bisected, with one half being used to display the cross-section to ensure that plating occurred evenly throughout the part, shown in Figure 7. The cross-section was ground with increasing fineness from 240 grit to 1500 grit and inspected under a stereo microscope with oblique illumination.115Figure 7. ENP cross-section and a micrograph of a single strut in the interior of the part prior to burnout.There are some macro defects visible in the plate (e.g., missing plating on the bottom at the third strut) that are due to the bisection with a band saw before the epoxy was embedded, causing the nickel plate to break. The minimum and maximum plate thicknesses measured were 760µm and 820µm.

The minimum and maximum plate thicknesses measured were 760µm and 820µm. The part plated consistently throughout, providing no evidence of a trend in plate thickness dependent upon the part geometry.When observed at a higher zoom with polarization under a different optical microscope, theplate shows a layering effect, visible in Figure 8.Figure 8. Polarized micrograph of a cross-section of the nickel plate showing plate layering with a SLS strut (black, top right) and an epoxy-filled pore (black, bottom left).The straight-lined crisscrossed pattern is due to the grinding process. Five layers (labeled) are readily visible, though more may become visible with a finer polishing procedure. It is believed that these layers are caused by the ENP replenishment process.

Five layers (labeled) are readily visible, though more may become visible with a finer polishing procedure. It is believed that these layers are caused by the ENP replenishment process. Extrapolating the width of the layers to include the range in which they are not visible in Figure 8 (moving from layer one to the nylon strut in the top right) provides a count of about 10 layers, the same as the number of replenishments during the plating process.116The uniformity between the layers shown indicates the consistency of the ENP process. The visible curvature is believed to be caused by an expected propagation of the surface roughness of the SLS part, leading to a part surface covered in small smooth nodules. This is better seen when closely viewing the surface of the part, shown in Figure 9.Figure 9.

This is better seen when closely viewing the surface of the part, shown in Figure 9.Figure 9. Micrograph of the surface finish of an ENP SLS part.Also present are irregularly located holes in the plate, seen as black circles on the nickel surface on the left side of Figure 9. These holes do not penetrate through the entire plate and are believed to be caused by the collection of gas bubbles formed during the ENP reaction, keeping the solution from reaching specific locations on the plate. This problem could likely be solved with ultrasonic agitation of the bath, although agitation was not recommended by the plating bath manufacturer (Caswell, Inc.).The second half of the bisected part moved to the burnout phase of the process with a flash burnout cycle in a gas fired kiln that heated from room temperature to 750°C in about 5 minutes, and then maintained temperature for 15 minutes. The furnace was then allowed to cool for about 20 minutes before being opened.

The furnace was then allowed to cool for about 20 minutes before being opened. The resultant complex, hollow, metal geometry is shown in Figure 10.Figure 10. Successful complex, hollow, metal geometry made with SLS, ENP and flash burnout (27mm wide).117The breaks in the plate toward the top of the part are due to the bisecting process and were present before burnout began. With a nickel plate thickness of about 800µm, the apparent density of this geometry is 3.16g/cm3 (less than half that of nickel at 8.9g/cm3), meaning that this structure is about 65 vol% air and 35 vol% nickel.5. CLOSURE AND FUTURE WORKThe authors presented work scoping and designing a hybrid manufacturing procedure for producing complex, hollow, metal geometries with the design freedom of AM systems. Any geometry that can be produced with SLS can undergo this process, opening a new class of geometries to designers.

Any geometry that can be produced with SLS can undergo this process, opening a new class of geometries to designers. Such geometries could be used for many applications, including novel heat exchangers and custom-fit high strength-to-weight ratio parts for structural supports or protective shielding.The authors identified and summarized a new manufacturing procedure for creating complex, hollow, metal geometries. They produced a complex nickel cellular structure that is 65 vol% air with 3mm pores and 800µm wall thicknesses. They used a polyamide substrate fabricated with selective laser sintering that underwent a three stage procedure of etching, room temperature metallization, and electroless nickel plating.The developed process has some limitations qualitatively determined during experimentation. Parts with pores smaller than 3mm may encounter difficulty during the RTM ENP process due to the large bubbles created that can become trapped in the part.

Parts with pores smaller than 3mm may encounter difficulty during the RTM ENP process due to the large bubbles created that can become trapped in the part. Since the part should not be agitated for the first few minutes of plating, the simple solution of agitation may not be possible. This also holds true for long channels that may hold air. The use of targeted low speed fluid pumps may help (perhaps running with an on/off duty cycle), but further investigation must be done to determine the associated detrimental effects to the start of plating.Future work will investigate the use of copper electroless processes to take advantage of the benefits of higher heat conductivity for heat exchanger applications. In addition to exploring new material systems, the authors look next to the fabrication of well-researched geometries (e.g., an octet-truss structure) in order to better compare the mechanical properties of parts.

Finally, the authors also look to the integration of part design with the process. Through collaborations with subject matter experts in fluid dynamics and heat transfer, the authors look to design a novel heat exchanger that can be compared to traditionally manufactured designs with standard metrics.6. ACKNOWLEDGMENTSThe authors acknowledge Dr. Carlos Suchicital, of the Kroehling Advanced Materials Foundry of Virginia Tech, for his help with cross-sectioning parts and burnout procedures. The authors also acknowledge Kevin Kline for his work calibrating and operating the Sinterstation 2500plus. David McCarthy was supported by the Department of Defense as part of the SMART Scholarship Program.1187. REFERENCES[1]D. T. Queheillalt and H. N. G. Wadley, "Pyramidal lattice truss structures with hollow trusses," Materials Science and Engineering: A, vol. 397, pp. 132-137, 2005.[2] M. Rombouts, J. P. Kruth, L. Froyen, and P.

57, pp. 996- 999, 2007.[21] R. Parkinson, "Plating on plastics. An industry review," Plating and Surface Finishing,vol. 81, pp. 32-35, 1994.[22] P. A. Schweitzer, Fundamentals of Corrosion: Mechanisms, Causes, and PreventativeMethods, 1 ed.: CRC Press, 2009.[23] M. Vasquez, B. Haworth, and N. Hopkinson, "Optimum sintering region for laser sintered nylon-12," Proceedings of the Institution of Mechanical Engineers. Part B, Journal of engineering manufacture, vol. 225, pp. 2240-2248, 2011.[24] D. L. McCarthy, "Creating Complex Hollow Metal Geometries Using Additive Manufacturing and Metal Plating," Master of Science, Mechanical Engineering, Virginia Polytechnic Institute and State University, 2012.120.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCreating Conformable Lithium Batteries Using Selective Laser SinteringT. Phillips1, C. Milroy2, J. Beaman11Department of Mechanical Engineering, University of Texas at Austin, TX 78712 2Texas Research Institute Austin, TX 78746AbstractSelective laser sintering is an additive manufacturing technique that uses a laser to consolidate powdered material and create complex three dimensional parts. Typically, SLS utilizes thermoplastic polymer media to create dense plastic components (direct SLS). It is also possible, however, to use composite powders with non-melting additives paired with a suitable binder to create highly functional materials (indirect SLS).

It is also possible, however, to use composite powders with non-melting additives paired with a suitable binder to create highly functional materials (indirect SLS). This paper will describe the formulation of composite materials containing conductive and electroactive material additives to fabricate lithium-ion battery components (i.e. anodes, cathodes, separators, and cases). Selective laser sintering adds to the geometric flexibility of the lithium battery components and enables batteries that conform to their surroundings, effectively reducing their geometric footprint. Preliminary galvanostatic charge/discharge test results will be presented for the functional Li-ion cathodes created using SLS, as well as next steps to improve capacity and reliability.IntroductionThe commercialization of lithium-ion (Li-ion) batteries in 1991 enabled the proliferation of miniaturized and mobile technology [1].

The high energy density, discharge rate, and lifetime have made Li-ion batteries the preferred method of energy storage for high-tech applications . Although rechargeable Li-ion batteries have powered mobile computing and telecommunications for decades, advances in energy-materials research has lagged behind global energy consumption and advances in microelectronics. As progress in improving the energy-density of insertion- compound-based electrodes has stalled, interest in “Beyond Lithium-ion” materials has increased.A frequently-overlooked opportunity to improve energy-density lies in the opportunity to optimize the form-factor of the battery to more efficiently integrate with devices to minimize footprint and reduce overall packaging needs.

By designing batteries for specific applications, the battery geometries can be tailored to fit around other components to reduce empty space and achieve greater energy storage without increasing device footprint.Additive manufacturing (AM) is a rapidly growing field for manufacturing complex parts with little-to-no tooling, making it an ideal technology for producing on-demand batteries with an improved footprint. One of the challenges to incorporating AM in battery manufacturing is the materials restrictions often faced by AM systems [2]. Li-ion batteries incorporate diverse metals (aluminum, copper, steel), inorganic energy-storage materials, and organic compounds (polymers, conductive carbons) that must act in synergy to create a functional energy-storage device.

Due to the high cost of AM equipment and material limitations, most published research efforts focus on a single battery component, typically active materials in the form of paste or slurry [3] [4].NAVAIR Public Release 2021-505 Distribution Statement A - 'Approved for public release; distribution is unlimited.’702Laser powder bed fusion (LPBF) has not been widely explored as a battery-making technology, yet offers diverse material options covering the materials needed for manufacturing of an entire battery; albeit with different machines within the LPBF family for building metal and polymer-based materials. LPBF as a battery manufacturing process introduces numerous advantages over traditional manufacturing techniques, such as facilitating creation of all battery components to create a truly conformable Li-ion battery, and substantially increasing the energy density by eliminating the current collector.

LPBF also presents some advantages over other AM techniques, such as the large build volume capable of creating nested and free-floating components, resulting in increased manufacturing speed [5].This paper will discuss progress on using indirect selective laser sintering (SLS) to create lithium battery components, with a focus on cathodes. Indirect SLS utilizes a multi-component feedstock, often with one melting component and one or more non-melting components [6]. The melting component acts as a binder for the non-melting components and post-processing steps must be taken to achieve the final desired material properties. In the work presented in this paper, a novolac phenolic resin powder was used as the melting component, which was subsequently crosslinked and pyrolyzed. Multiple material compositions have been studied and preliminary charge and discharge data will be presented for one such composition.

Multiple material compositions have been studied and preliminary charge and discharge data will be presented for one such composition. To iterate material compositions more quickly, reduce wasted material, and optimize electrode performance, only planar, single-layer electrodes were considered.Lithium Ion BatteriesTo understand the challenges of using AM to produce batteries, the basics of Li-ion battery (LiB) operation must be understood. A lithium ion battery provides controlled electrochemical energy storage and release through the cooperative functionality of 6 major material components: cathode, anode, separator, electrolyte, current collector, and case. In commercial LiB, the anode and cathode are based on insertion materials that intercalate lithium ions. The LiB anode, or negative electrode, contains graphite, which has a very low thermodynamic reduction potential (Eo) close to that of lithium metal.

The LiB anode, or negative electrode, contains graphite, which has a very low thermodynamic reduction potential (Eo) close to that of lithium metal. The graphite hosts lithium ions between its layered sheets when the battery is in the charged state. In contrast, the cathode, or positive electrode, contains electrochemically active material(s) with a high thermodynamic reduction potential, typically in the form of a lithiated transition metal (Mn, Fe, Co, Ni) oxide or phosphate. The difference between the thermodynamic reduction potentials of the anode and cathode determines the operational voltage of the battery, and causes electrons to flow spontaneously (ΔG < 0) from anode to cathode (galvanic discharge) while lithium ions migrate from anode to cathode to compensate the electrostatic charge accumulation in the cathode from the flow of electrons.

In this way, the battery can provide electrical current to an external circuit during discharge, but will later require electronic current from a power supply to recharge the battery. Unfortunately, most electroactive materials used in batteries are electronic semiconductors or insulators, so commercial LiB cathodes typically contain carbon additives to increase electrical conductivity, and a polymer binder to secure the active material and conductive carbon to a metal current collector. The electrodes are kept from touching each other with a thin polymer or ceramic separator that is electronically insulating (to prevent electrical shorting and fires) but ionically conductive (to permit the passageNAVAIR Public Release 2021-505 Distribution Statement A - 'Approved for public release; distribution is unlimited.’703of lithium ions from the anode to cathode during discharge, and from cathode to anode during charge).

The components are flooded with an organic carbonate electrolyte solution containing lithium salts that facilitate lithium transfer between the electrodes. The anode and cathode are mounted on electrically-conductive current collectors that provide electrical contact between the battery and an external circuit. All these components are housed within a case that protects the components and keeps the electrolyte contained. The lithium ions shuttle from the anode to the cathode during discharge (a galvanic process), and return to the anode during charging, under an electrical load from an external charger. This reversible shuttling two insertion/intercalation materials is often referred to as the “rocking chair’ operational principle.The current process for commercial LiB manufacturing involves multiple steps in which a slurry is first prepared and then coated and pressed on to metal foils (copper for anodes, aluminum for cathodes) using large roll-to-roll scale tooling.

The electrodes are subsequently cut then and rolled (for cylindrical cells) or folded (for prismatic or pouch cells), with a separator sheet between each individual anode and cathode sheet. After assembly, the cells are filled with electrolyte and sealed.Materials and MethodsLi-ion battery cathodes were created in an SLS system (Sinterstation HiQ) using a composite material system containing phenolic, carbon fibers, graphite, and lithium iron phosphate (LFP). The carbon fibers (Zoltek PX-30) had a mean length of 100 µm and width of 7 µm and composed 36.9 weight percent of the mixture. The LFP (MSE Supplies) had a mean particle diameter of 1.5 µm and composed 43.1 weight percent of the mixture. The graphite (Imerys C- nergy SFG 15L) had a mean particle diameter of 8.8 µm and composed 5 weight percent of the mixture. The remaining 15 weight percent consisted of cryo-milled phenolic (Georgia Pacific 5520).

The remaining 15 weight percent consisted of cryo-milled phenolic (Georgia Pacific 5520). Electrodes were 12 mm diameter circles whose thickness was controlled by the laser melt penetration.Electrode processing in the Sinterstation used a nitrogen atmosphere with less than 0.1 % oxygen and a preheat temperature of 65 ⁰C. Laser power was set to 20 watts and scanned at the standard speed, presumably 10,160 mm/s. Because of the large quantity of material required for a multilayer build in the Sinterstation, a pseudo multilayer technique was adopted. A thick layer (approximately 5 mm) of powder was spread, and the electrode cross sections were rescanned multiple times. Each successive scan reheats the electrode cross section and increases the depth of penetration. Forty scans were needed to penetrate deep enough into the powder bed to create an electrode with suitable strength and thickness.

Forty scans were needed to penetrate deep enough into the powder bed to create an electrode with suitable strength and thickness. Successive scan lines (hatch spacing) were spaced 0.254 mm apart.After removal from the SLS system, cathodes were crosslinked and pyrolyzed (thermally decomposed). Crosslinking was performed by placing electrodes in a preheated oven at 175 °C for 2 hours. Preheating the oven allowed the phenolic to quickly reach crosslinking temp, minimizing the amount of polymer reflow and geometric distortion experienced by the electrodes. After crosslinking, the electrodes were pyrolyzed in a vacuum furnace at approximately 300 mTorr.

After crosslinking, the electrodes were pyrolyzed in a vacuum furnace at approximately 300 mTorr. The temperature was ramped at 3 ⁰C/min to 800 °C and held for 4 hours before cooling naturally.NAVAIR Public Release 2021-505 Distribution Statement A - 'Approved for public release; distribution is unlimited.’704Scanning electron microscopy (SEM) and X-ray diffraction (XRD) were used to characterize the electrodes at different stages of post-processing.Electrodes were stored in vacuum desiccators between each processing step and prior to electrochemical testing.

Galvanostatic charge-discharge testing was carried out in coin-cell format and half-cell configuration with lithium metal foil anodes, Celgard 2500 polypropylene separators, nickel-foam spacers, and an electrolyte comprising lithium hexafluorophosphate dissolved in ethylene carbonate and ethyl-methyl carbonate (1 M LiPF6 in 1:1 EC:EMC).ResultsXRD spectra for the individual components (graphite, carbon fiber, phenolic, and LFP) and the resulting electrode are given in Figure 1. Spectra for all constituent materials were taken in the material’s “green”, or as-purchased, state while the spectrum for the final electrode was after SLS processing and crosslinking.

Spectra for all constituent materials were taken in the material’s “green”, or as-purchased, state while the spectrum for the final electrode was after SLS processing and crosslinking. Electrodes in the green, crosslinked, and pyrolyzed states were able to be handled but were fairly weak and would fracture if dropped onto a hard surface from a height of 50 cm.Figure 1: Normalized X-ray diffraction spectra for graphite, carbon fiber, phenolic, lithium iron phosphate, and an additively manufactured battery electrode containing all four materials.SEM images of the raw materials used to additively manufacture a lithium-ion battery cathode are given in Figure 2.

EDS analyses are included in appendix A.NAVAIR Public Release 2021-505 Distribution Statement A - 'Approved for public release; distribution is unlimited.’705ABCDFigure 2: SEM images of battery material used for additive manufacturing of lithium batteries: (A) Carbon fibers, (B) Cryo-milled phenolic powder, (C) Lithium iron phosphate, and (D) GraphiteSEM of the as-built electrode shows significant agglomeration of LFP and material heterogeneity. Figure 3 offers images of a crosslinked cathode at different magnifications. In Figure 3(A) cracks can be seen throughout the electrode, contributing to its poor mechanical strength. Material heterogeneity can also be seen, as evidenced by the large portions or phenolic-rich material at the center of Figure 3(B) surrounded by LFP-rich regions. LFP agglomerates canNAVAIR Public Release 2021-505 Distribution Statement A - 'Approved for public release; distribution is unlimited.’706be clearly seen in Figure 3(C).

LFP agglomerates canNAVAIR Public Release 2021-505 Distribution Statement A - 'Approved for public release; distribution is unlimited.’706be clearly seen in Figure 3(C). It is suspected that the poor electrical contact between LFP particles impedes electrical or ionic transport to the interior particles of the agglomerates and the poor distribution of phenolic binder leads to weaker electrodes that have a decreased lifespan than those produced using traditional techniques. Energy-dispersive X-ray spectroscopy (EDS) analysis of a pyrolyzed cathode is given in Figure 4, showing the distribution of material in the electrode.ABCFigure 3: SEM images of a crosslinked cathode at different magnifications. At low magnification (A), cracks in the electrode are evident, contributing to poor mechanical strength. At medium magnification (B), the individual materials can be seen and material heterogeneity is clear.

At medium magnification (B), the individual materials can be seen and material heterogeneity is clear. At high magnification (C), individual LFP particles as well as many LFP agglomerates can be seen.NAVAIR Public Release 2021-505 Distribution Statement A - 'Approved for public release; distribution is unlimited.’707Figure 4: EDS analysis of pyrolyzed electrode showing the material distribution.Figure 5 presents a representative plot of extended galvanostatic charge-discharge cycling of a lithium iron phosphate (LFP) cathode for one-hundred cycles at C/10 discharge rate. The cathode initially produces nearly one-third of the theoretical capacity of LFP (170 mAh/g), and the capacity and coulombic efficiency (CE) were somewhat erratic during the first five cycles. Thereafter, cycle performance was steadier, but the cathode suffered mild yet constant capacity fade.

Thereafter, cycle performance was steadier, but the cathode suffered mild yet constant capacity fade. The coulombic efficiency was initially low, which is hallmark of Li-ion cells during formation of the solid-electrolyte interface (SEI) that is essential for stable long-term cycle performance. Although the CE remained above 98% after cycle ~ 12, it was still lower than a typical commercial LiB.Figure 5: Extended galvanostatic charge/discharge testing of a cathode produced via additive manufacturing.NAVAIR Public Release 2021-505 Distribution Statement A - 'Approved for public release; distribution is unlimited.’708Conclusion and DiscussionIndirect selective laser sintering using phenolic binder was evaluated for manufacturing Lithium-ion battery electrodes. Cathodes containing Lithium iron phosphate, carbon fibers , graphite, and phenolic were additively manufactured and post-processed using a high-temperature vacuum furnace.

Cathodes containing Lithium iron phosphate, carbon fibers , graphite, and phenolic were additively manufactured and post-processed using a high-temperature vacuum furnace. Half-cell testing showed the cathode had nearly one third the theoretical capacity of LFP but experienced significant capacity fade through the first 100 cycles. While underperforming compared with traditionally manufactured lithium-ion batteries, additive manufacturing provides manufacturing flexibility that can enable on-demand manufacturing of application-specific batteries.The results presented in this paper indicate indirect SLS is a viable method of manufacturing Li-ion battery components. Continuing work will focus on improving energy capacity and lifecycle of the electrodes. SEM analysis revealed a high amount of agglomeration and heterogeneity in the powder feedstock, both of which can contribute to decreased battery performance.

SEM analysis revealed a high amount of agglomeration and heterogeneity in the powder feedstock, both of which can contribute to decreased battery performance. Ball milling techniques will be explored to break up agglomerates and make the LFP material more electrically available. Additional mixing methods will be explored to improve the homogeneity of the material mixture to better distribute the phenolic binder.An additional path of interest for future work is to evaluate phenolic’s role in the battery and how it is affected by post-processing. Phenolic is a suitable binder in this system due to it being a high-ash polymer, converting to conductive carbon upon pyrolysis. Other researchers have found that the conversion and conductivity is influenced by pyrolysis time, temperature, and vacuum level [7] [8] [9].

Other researchers have found that the conversion and conductivity is influenced by pyrolysis time, temperature, and vacuum level [7] [8] [9]. Adjusting the post-processing steps to maximize electrical conductivity while maintaining structural integrity could improve the battery performance of AM electrodes.FundingThis project is supported by Phase II funding from NAVAIR STTR topic N18A-T008 (Additive Manufacturing for Naval Aircraft Battery Applications), Contract number N68335-19- C-0578. The authors gratefully acknowledge Dr. Adam Jolley and Michael Melnick (technical points of contact at NAVAIR) for their guidance and support.NAVAIR Public Release 2021-505 Distribution Statement A - 'Approved for public release; distribution is unlimited.’709Works Cited[1] J. T. Warner, Lithium-Ion Battery Chemistries: A Primer, Elseview, 2019.[2] D. Bourell, J. P. Kruth, M. Leu, G. Levy, D. Rosen, A. M. Beese and A. Clare, "Materials foradditive manufacturing," CIRL Annals, vol.

EATING EMBEDDED RADIOFREQUENCY STRUCTURES USING POLYJET MATERIAL JETTINGRichard L. Dumene, Paul Kennedy, Christopher B. Williams1, Dennis Sweeney2, Greg Earle3 1 Design, Research, and Education for Additive Manufacturing Systems Laboratory Department of Mechanical Engineering, Virginia Tech 2Director of Instructional Laboratories Bradley Department of Electrical and Computer Engineering 3 Center for Space Science and Engineering Research Bradley Department of Electrical and Computer EngineeringABSTRACTEmbedding of integrated systems via Additive Manufacturing (AM) offers the potential to save weight, space, and time in the production of electronics and vehicles. Of specific interest are embedded electrical systems that operate in the radiofrequency (RF) range as they have great potential utility in communication systems and aircraft avionics including navigation.

However, systems in this frequency range pose unique manufacturing challenges such as the need to minimize impedance discontinuities in the system. In this work, the authors explore various techniques for embedding RF structures such as antennas and filters via a multi-material jetting AM process. Specifically, the dielectric constants and loss tangents of Vero White and Tango Black polymers were determined to facilitate the design of RF structures with these materials. It is shown that measurements of S-parameters of the resultant embedded RF structures approach or match the performance of non-embedded electronics.1. INTRODUCTION1.1.

It is shown that measurements of S-parameters of the resultant embedded RF structures approach or match the performance of non-embedded electronics.1. INTRODUCTION1.1. Embedding Conductive Materials into Additively Manufactured PartsDue to its ability to provide access to the entire volume of a part throughout the build process (Kumar, 1998), Additive Manufacturing (AM) is uniquely positioned to provide designers the opportunity to realize multifunctional parts by embedding objects into the parts as they are fabricated. In general, embedding foreign components in an AM build process involves (a) designing a part with a specifically designed void to accept the foreign component, (b) pausing the print mid-build, (c) inserting the foreign object into the designed void, and (d) resuming the print over the inserted component.

Past examples of embedding objects in Additive Manufacturing processes include embedded motors via Stereo lithography (SL) (De Laurentis et al., 2002), optical sensors via Shape Deposition Manufacturing (Li and Prinz, 2003), and shape memory alloys (Meisel et al., 2014), LEDs and IR components in PolyJet material jetting (Willis et al., 2012).Recent research has been focused in directly depositing conductive materials into printed layers in order to create objects that feature embedded circuitry, sensors, and/or antennae. This capability is of great interest in integrated electronics manufacturing due to its potential to save weight, space, and time in the production of electronic components through the embedding of integrated systems into a printed part’s structure.

This capability is also beneficial to AM ingeneral, as it provides an opportunity to integrate the manufacturing of multi-material complex systems into a singular manufacturing process.The majority of effort in printing components with embedded circuitry has focused in the integration of Direct Write (DW) and AM technologies. DW is a set of technologies that enable selective deposition of material onto a substrate, and include extrusion, droplet jetting, aerosol jetting, laser transfer, and tip transfer technologies (Hon et al., 2008). The integration of DW and AM technologies the creation of complex and conformal electronics (interconnects, sensors, etc.) that are structurally integrated into a manufactured part.

The integration of DW and AM technologies the creation of complex and conformal electronics (interconnects, sensors, etc.) that are structurally integrated into a manufactured part. One of the first examples of integrating the two technologies was demonstrated by a research team at the University of Texas El Paso (UTEP); the hybrid process worked by first pausing the SL build, using a direct write device to extrude a conductive paste onto the layer in the pattern of the desired circuit, and then resuming the build to embed the printed circuit (Lopes et al., 2005; Medina et al., 2005; Palmer, 2005).

Additional examples of hybridizing DW and AM to create parts with embedded circuits have also been demonstrated in Ultrasonic Consolidation (Robinson et al., 2006), screw extrusion (Malone et al., 2008), fused filament extrusion (Breyfogle and Cormier, 2013), polymer powder bed fusion (Folgar et al., 2014), and multi-material jetting processes (Perez and Williams, 2014a).Following an in-depth analysis of the integration of DW and AM technologies, it was observed that the direct application of dissimilar materials is challenging as it is hampered by (i) identifying DW materials and processes that are compatible with AM infrastructure, throughput and resolution, (ii) temperature processing requirements, and (iii) interactions between the two materials (Perez and Williams, 2014b).

These physical limitations have resulted in embedded circuits that have fairly high resistivity, which limits their application.Due to these challenges, many are investigating other means of incorporating conductive materials into printed layers. Specifically, instead of directly writing conductive materials onto a printed substrate, researchers are investigating the embedding of conductive copper wire and direct placement of pre-fabricated sensors and integrated circuits. Early efforts in this approach were focused in embedding wires and sensors into machined pockets in Ultrasonic Consolidation builds (Siggard et al., 2006). More recent research from the UTEP group has demonstrated the use of an ultrasonic wire embedding device to directly embed/write copper wire into polycarbonate printed via Extrusion AM (Kim et al., 2015).

This system has enabled the creation of complex systems such as a CubeSat featuring embedded interconnects that are laser-welded to surface mount devices (Espalin et al., 2014). Copper wire and coper mesh has also been embedded into polycarbonate material to create capacitive sensors (Shemelya et al., 2015).1.2. Fabricating Radio Frequency Structures via Additively ManufacturingAs noted above, most of the research in embedding electronics into AM components has focused on direct printing sensors and interconnects. There is emerging interest in fabricated embedded electrical systems that operate in the radiofrequency (RF) range, as they have great utility in communication systems and aircraft avionics, such as radar. Highly integrated systems with limited space such as CubeSATs and small UAVs would significantly benefit if antennas and receivers were able to be embedded in the structure of the CubeSAT.

Highly integrated systems with limited space such as CubeSATs and small UAVs would significantly benefit if antennas and receivers were able to be embedded in the structure of the CubeSAT. Similarly, the design of 3D printed GHz antennas could one day facilitate a 3D printed cell phone. Another possibility is to embed radar hardware in the wings of UAVs. In order to realize these possibilities, certainchallenges in the manufacture of these systems must be overcome including the need to minimize impedance discontinuities in the system and the need to characterize the electromagnetic properties of materials used in AM.The University of Texas at El Paso research groups has conducted research in the area of embedding RF structures into AM parts. Deffenbaugh used stereo lithography (SL) and fused deposition modeling (FDM) to build RF structures (2014). Specifically, SL was used to build micro strip transmission lines, a strip line, and a suspended micro strip structure.

Specifically, SL was used to build micro strip transmission lines, a strip line, and a suspended micro strip structure. FDM methods were used to build a curve inverted-F antenna. This research demonstrated that RF structures can be constructed using additive manufacturing technologies. Deffenbaugh identified two key obstacles for process development, (i) the FDM system has a relatively low-resolution and (ii) SL is limited in its embedding abilities as it relies on curing resin in a vat. Similarly, a patch antenna was created by Robinson and coauthors by using a hybridized DW and Ultrasonic Consolidation system (2006).1.3. ContextThe goal of this work is to explore the use of the PolyJet material jetting process as a means of fabricating printed parts with embedded RF structures. The primary objects are to demonstrate the feasibility of embedding RF structures in this process, and to assess their resultant quality (via S-parameter measurements).

The primary objects are to demonstrate the feasibility of embedding RF structures in this process, and to assess their resultant quality (via S-parameter measurements). As discussed in Section 2, the PolyJet process has both unique advantages (e.g., the ability to fabricate multi-material components that have tailored stiffness and flexibility) and unique disadvantages (e.g., the low heat deflection temperature of the printed polymers), that make this a challenging problem. Through the developed embedding technique, the authors successfully fabricated five different RF structures that are embedded in three different PolyJet material systems. Details of the design of these structures are presented in Section 3, along with a description the experimental methods employed to analytically predict and experimentally measure their performance. The results of the fabrication process and the measurements of the RF structures’ performance are presented in Section 4.

The results of the fabrication process and the measurements of the RF structures’ performance are presented in Section 4. Closure is offered in Section 5.2. EMBEDDING RF STRUCTURES IN POLYJET MATERIALSPolyJet is a material jetting AM process that uses drop-on-demand inkjet printing to selectively deposit droplets of photopolymer directly onto a build platform. Once a layer of droplets is deposited via inkjet print heads, a roller evens out the layer surface, and two UV lights (one leading and one trailing) pass over the printed layer multiple times to fully cure the photopolymer. As several inkjet print heads with separate material sources can be installed into the printing block, multiple materials can be deposited in a single layer, thus enabling the creation of graded materials. For example, the use of Vero White Plus (a stiff polymer) and Tango Black Plus (a flexible, elastomeric material) enables a designer to create products with graded stiffness.

For example, the use of Vero White Plus (a stiff polymer) and Tango Black Plus (a flexible, elastomeric material) enables a designer to create products with graded stiffness. A sacrificial support material is also deposited during a build to support the fabrication of complex geometries. This sacrificial material is removed using a compressed water jet in a post-processing step.The PolyJet process creates new layers by direct material addition (Williams et al., 2011), which is well suited for component embedding. The lack of a recoating step, wherein raw material is added in bulk in powder or resin form, eliminates concerns of disturbing the previously printed layer typically found in embedding processes.

The lack of a recoating step, wherein raw material is added in bulk in powder or resin form, eliminates concerns of disturbing the previously printed layer typically found in embedding processes. The authors have demonstrated the use of PolyJetto realize products with embedded monofilament (Stiltner et al., 2011), shape memory alloy actuators (Meisel et al., 2014), direct write circuits (Perez et al., 2014a), and other photopolymer materials (e.g., quantum dot composites resins (Elliott et al., 2013)). Although different in form and function, each of these components was embedded into a part during printing by following a generalized four-step process: (i) designing a cavity for the component, (ii) pausing the build at the topmost layer of the cavity, (iii) inserting the component and (iv) resuming the build.

One of the key aspects of the embedding process for PolyJet is the need to remove support material from the cavities in the printed part in order to properly place any desired embedded components. In addition, since the print head assembly block passes over the printed part at a clearance of only 100 µm, embedded components cannot protrude from the previously deposited layer, or damage to the printing nozzles could occur.The authors have previously embedded electronic components in PolyJet builds by inserting a pneumatically-controlled extrusion printer into the PolyJet system, and selectively depositing a conductive paste onto a printed layer (Perez et al., 201b). While this was successfully used to deposit interconnects and to fabricate a strain gauge, the low heat deflection temperature of the printed photopolymer (54°C) prevented the application of the recommended heat treatment for the deposited material.

Thus, the deposited conductive traces had high resistivity, and prevented the creation of electronic structures that required significant current.To address this limitation, the authors look to the direct embedding of copper materials to create electronic components, as done previously in Extrusion AM (Section 1.1). However, instead of embedding copper wire or mesh (Espalin et al., 2014; Shemelya et al., 2015), the authors look to the use of copper foil adhesive tape as a means of creating RF structures. Copper tape is conventionally used for creating low-profile conductive traces, antenna fabrication, and for RF- shielding. It is sufficiently conductive for RF applications, and its width enables the creation of designed filters and antennae structures by precision cutting (via blade or laser cutter) into a designed shape. The adhesive backing also allows for efficient application and anchoring into a pre-designed cavity.

The adhesive backing also allows for efficient application and anchoring into a pre-designed cavity. This approach proved to be much faster and cheaper than selectively extruding nano-silver conductive paste.To facilitate the embedding of copper tape, the designed part must be altered to feature a sufficient cavity to accept the tape in order to prevent collision of the objet with the print heads. When paused, support material is removed using precision tools; the exposed surface of the part is then cleaned using 99% isopropyl alcohol to remove residual support material and prepare the surface for adhesives. Additionally, access holes for tooling access into the part must be designed and produced. The design of access holes is particularly challenging for ensuring access for the soldering iron, which is used to attach connectors to the embedded RF structures.

The design of access holes is particularly challenging for ensuring access for the soldering iron, which is used to attach connectors to the embedded RF structures. The soldering iron must not be permitted to touch the material as the melting point of the cured photopolymers is significantly lower than that of solder (SAC350). The most successful approach required the use of wide, shallow holes to facilitate easy access with a soldering iron. After the connectors are affixed to the conductors, any access windows or ports can be filled with epoxy or photopolymer and then cured via UV.3. EXPERIMENTAL METHODSThe fabricated RF structure, and the experimental methods used to measure them, are presented in this section. All samples were printed using a Stratasys Connex 350 PolyJet AM system and were made from either Vero White Plus (VW), Tango Black Plus (TB), or Shore95, a composite of the VW and TB materials.3.1.

Characterizing Dielectric ConstantsThe dielectric constant of the substrate material determines the dimensions of the transmission line structures. Initial estimates of the dielectric constant were produced based on the impedance of early attempts at building transmission lines. To more accurately determine the dielectric constant, capacitor test fixtures were constructed. Capacitor structures were chosen because capacitors are well understood and their capacitance is a clear function of dielectric constant. The procedure for determining the dielectric constant first involved printing 1” square capacitors from the 3D printed materials (VW, TB, and Shore95) to serve as substrates from copper tape, which served as the plates. Once the embedding and printing was complete, the authors measured the capacitance and then calculated the dielectric constant.Figure 1.

Once the embedding and printing was complete, the authors measured the capacitance and then calculated the dielectric constant.Figure 1. Q-meter and equivalent circuit schematic.The capacitance measurements were obtained via a 4342A Hewlett Packard Q-meter (Figure 1). The Q-meter circuit is an inductor in parallel with a variable capacitor. The variable capacitor was adjusted until the circuit reached resonance. Then, the capacitor under test was placed in parallel with the variable capacitor. The variable capacitor was again adjusted to tune the circuit. The difference between the first and second values of the variable capacitor was equivalent to the capacitance of the capacitor under test. This capacitance value, along with the dimensions of the capacitor, was used to calculated the dielectric constant, εr.3.2. Design and Fabrication of RF StructuresFollowing the process outlined in Section 2, several embedded RF structures were made using VW, TB, and Shore95 materials.

Design and Fabrication of RF StructuresFollowing the process outlined in Section 2, several embedded RF structures were made using VW, TB, and Shore95 materials. The authors chose well understood structures so that well known analytical models could be used to design the dimensions for these structures, and so that commercial simulation software could be used to verify that the fabricated structures were performing appropriately (Section 3.4).Coaxial cable: The first embed consisted a simple copper clad coaxial cable. This sample was a 0.047 inch copper cladded coaxial cable embedded in an 8 inch (20cm) strip of Vero White photopolymer. For simplicity, this part was embedded without connectors using a simple U channel. After embedding the cable, it was secured to the structure using SuperGlue before resuming the print. This prevented the part from rising off of the build platform and potentially damaging the print heads.

This prevented the part from rising off of the build platform and potentially damaging the print heads. After successfully embedding the cable, SMA connectors were soldered to the cable.Micro strip transmission lines: A micro strip was made from VW, and was designed to have an impedance of 50 ohms. Copper tape of different widths was affixed to a substrate 0.1” in height. The connectors were then soldered to the copper tape. Two micro strip transmission lines were created; one mounted on the surface of the printed substrate, and one completely encapsulated by the printed material.Low-pass filters: A third order low pass filter was created by embedding copper tape onto printed VW and Shore95. The filter design consisted of two sections of shunt capacitance and a central inductive section. The low pass filters were designed to be Butterworth filters with a -3 dB point at 1.6 GHz and a filter impedance of 50 ohms.

The low pass filters were designed to be Butterworth filters with a -3 dB point at 1.6 GHz and a filter impedance of 50 ohms. The filters were designed to have three elements that would behave like the equivalent circuit shown in Figure 2. In order to build this circuit in a micro strip line, a high low impedance section of 25 ohms was followed by a high impedance section of 100 ohms, and then another 25 ohm low impedance section. The filters were connected to 50 ohm SMA connectors by 50 ohm micro strip lines. The substrate heights were chosen such that a micro strip width of 0.25” would produce a characteristic impedance of 50 ohms.Figure 2. Equivalent circuit for designed low pass filterFigure 3. Equivalent circuit for designed edge coupled single resonator band pass filterEdge coupled single resonator band pass filter. Edge-coupled micro strip band pass filters consist of micro strip feeds capacitively coupled through micro strip resonators.

Edge-coupled micro strip band pass filters consist of micro strip feeds capacitively coupled through micro strip resonators. The band pass filter in this work used utilized a single half-wavelength resonator. The equivalent circuit for a single resonator edge-coupled micro strip band pass filter is shown in Figure 3. The filter was designed to have a center frequency of 1.58 GHz. Accordingly, the resonator was designed to be 61.645 mm in length. The substrate height was chosen to be 2.54 mm due to the dimensions of the SMA connectors. Accordingly, the width of all the micro strips was calculated to be 7.13 mm in order to produce a micro strip line with an impedance of 50 ohms. The spacing between the resonator and the micro strips was determined by iterative simulations using QUCs simulation software.

The spacing between the resonator and the micro strips was determined by iterative simulations using QUCs simulation software. A spacing of 1 mm between micro strips was chosen based on manufacturing tolerances, minimal return loss at the center frequency in the QUCs simulation, and a bandwidth of 80 MHz in the QUCs simulation.Micro strip patch antennas. Micro strip patch antennas with micro strip feeds were fabricated to explore the potential for producing 3D printed antennas. Two patch antennas with the same dimensions were produced on Vero White and Shore95 substrates. The antennas were designed to work at approximately 1.575 GHz in order to be able to receive GPS signals. The height of the substrate was chosen such that a 6.35 mm micro strip would have a characteristic impedance of 50 ohms. To compensate for the effects of fringing fields, the calculated dimensions were optimized based using FEKO simulations to maximize poweroutput at 1.575 GHz.

To compensate for the effects of fringing fields, the calculated dimensions were optimized based using FEKO simulations to maximize poweroutput at 1.575 GHz. Given the widespread use of GPS in cell phones, UAVs, and CubeSATs building a 3D printed GPS antenna advances the progress towards developing a fully integrated 3D printed UAV or cell phone.3.3. Characterizing RF PerformanceA Tektronix 1502 Time Domain Reflectometer (TDR) Cable Tester was used to determine the location of discontinuities and impedance changes within the fabricated structures. The TDR sends an impulse down the device and records the time delay before the reflection reaches the TDR. It then calculates the reflection coefficient as a function of length based on this time delay. Any deviations in the reflection coefficient signifies an impedance mismatch or discontinuity within the transmission line.

Any deviations in the reflection coefficient signifies an impedance mismatch or discontinuity within the transmission line. Since the TDR provides this information as a function of length, the discontinuity can be easily located on the device under test. This device was used to determine the quality of solder joints between the SMA connectors and the transmission lines, and also enabled the comparison and visualization of the impedance of multiple sections of an embedded transmission line.The performance of the radiofrequency structures was characterized by a two port measurement of the scattering (S-parameter) matrix using 2 port S-parameter measurements on a Rohde & Schwarz ZVL13 Network Analyzer. The network analyzer was designed to work from 9 kHz to 13.5 GHz. The impedance, as a function of frequency, was also calculated from the reflection coefficient, S11. The devices were characterized over a frequency range from 0.1 to 10 GHz.

The impedance, as a function of frequency, was also calculated from the reflection coefficient, S11. The devices were characterized over a frequency range from 0.1 to 10 GHz. In order to ensure the accuracy of the measurements, the network analyzer was calibrated before each measurement session using a ZV-Z135 Rohde & Schwarz calibration kit.When analyzing the S-parameter measurements, an S21 measurement of 0 dB represented a perfect transmission line. Similarly, an S11 measurement of 0 dB represented a perfect open or short. When analyzing filters, the cutoff frequency, or -3 dB point, occurs when the S21 parameter is 3 dB below its maximum point. This cutoff frequency is significant as it represent where half of the maximum power is transmitted.3.4. Simulation of Fabricated RF StructuresThe authors employed two separate simulation software packages to verify that the designed RF structures were performing as expected.

Simulation of Fabricated RF StructuresThe authors employed two separate simulation software packages to verify that the designed RF structures were performing as expected. First, Qucs (Quite Universal Circuit Simulator) was used to build schematic representations of the samples and predict their S-parameter values with changes in frequency. Qucs is a circuit simulator similar to Agilent ADS in that it can simulate micro strip and strip line circuits, including filter geometries, in the GHz range. Major advantages of Qucs included the short simulation time and the short time required to construct schematics for simple structures.Qucs, however, does not support producing 3D models of the structures nor simulation of antenna designs. As two of the major goals of this project were to utilize the complex geometries possible with additive manufacturing and to design antennas for embedding, another software platform was needed.

As two of the major goals of this project were to utilize the complex geometries possible with additive manufacturing and to design antennas for embedding, another software platform was needed. FEKO, an EM Simulation Software, was used to build 3D representations of the structures, and then to simulate the S-parameters, impedances, and far fields of the various structures as a function of frequency. While simulating structures in FEKO is considerably moretime intensive, the simulation results in FEKO were more accurate than the results from Qucs. The difference in performance became more significant as frequency increased. The deviation in performance of these two modelling tools can be seen in the discussion of the Vero White low- pass micro strip filter specimen (Section 4.4). Ultimately, both tools showed potential as tools in the design of 3D printed RF structures.4. RESULTS AND DISCUSSION4.1.

Ultimately, both tools showed potential as tools in the design of 3D printed RF structures.4. RESULTS AND DISCUSSION4.1. Dielectric Constants of PolyJet MaterialThe average dielectric constants and 95% confidence intervals for VW, TB, and Shore95 (a composite of VW and TB) materials are shown in Table 1. The interval for VW and Shore95 are small, but the range for TB is wider. While this range could be due to variations in the printed material, it is likely a result of challenges in measuring the capacitance of the TB capacitors. TB is a flexible material, so the capacitors would bend, causing the copper tape to peel off of the TB dielectric.Table 1.

TB is a flexible material, so the capacitors would bend, causing the copper tape to peel off of the TB dielectric.Table 1. Dielectric Constants and 95% Confidence IntervalsMaterialAverage Dielectric Constant95% Confidence IntervalVero White2.9521[2.9082, 2.9960]Tango Black2.8161[2.4563, 3.1758]Shore953.1936[3.1245, 3.2628]Initially, the empirical results of the structures did not match expectations based on simulations of the structures. This discrepancy was due to the notable loss tangent of the VW, which was used in most of the initial structures. A loss tangent of 0.0285 was calculated for the VW substrate by fitting the simulated curve to the empirical results. For comparison, the loss tangents of VW and other common materials are shown in Table 2. VW is far more lossy than industry standard FR- 4, and more closely resembles the large loss of wood.

For comparison, the loss tangents of VW and other common materials are shown in Table 2. VW is far more lossy than industry standard FR- 4, and more closely resembles the large loss of wood. This large loss tangent suggests there are limitations to the utility of Vero White substrates in the construction of transmission lines. However, the potential exists to utilize the complex geometries possible with AM to create VW structures with less loss.Table 2. Common Materials and Dielectric CharacteristicsMaterialRelative PermittivityLoss TangentFused quartz3.80.00006 @ 3 GHzTeflon (PTFE)2.0 - 2.10.00028 @ 3 GHzFR-4 -high resin4.20.008 @ 3 GHzABS (plastic), Molded2.0 - 3.50.00500 - 0.0190Vero White2.970.0285 @ 1.5 GHzWood1.2 - 2.10.03 @ 3 GHz4.2. Coaxial CableAs presented in Section 3.2, the first embedded object was a copper cladded coaxial cableembedded into a strip of VW (Figure 4).Figure 4.

Coaxial CableAs presented in Section 3.2, the first embedded object was a copper cladded coaxial cableembedded into a strip of VW (Figure 4).Figure 4. An early embedded coaxial cable demonstrating embedding capabilities.Following the soldering of SMA connectors, a network analyzer was used to evaluate the system. This system performed well into the GHz range as expected. It can be seen in Figure 6 that the S21 measurement was almost 0 dB across the entire spectrum indicating a near perfect transmission line. Thus, GHz structures can be created using simple embedding techniques and widely available materials such as coaxial cables. Using coaxial cable isolated the RF performance of the TEM structure from the tolerances and errors inherent in the printing process.

Using coaxial cable isolated the RF performance of the TEM structure from the tolerances and errors inherent in the printing process. This type of system would be most useful for connecting application specific sections, such as filters and antennas, when the transmission path may have to move in three dimensions or outside of a printed part.Figure 5. Parameter measurements from the embedded coaxial cable. The S21 of almost 0 dB indicates an almost ideal transmission line.4.3. Micro strip transmission lineThe initial micro strip transmission line (Figure 6) performed well below 5 GHz and had an impedance of approximately 55 ohms (Figure 7), as calculated from the S11 measurements (Section 3.2). However, the copper tape would kink when the part was moved, causing variations in the impedance visible on the TDR.Figure 6. An embedded micro strip transmission lineFigure 7.

However, the copper tape would kink when the part was moved, causing variations in the impedance visible on the TDR.Figure 6. An embedded micro strip transmission lineFigure 7. S-Parameter results for the micro strip transmission line showed deteriorating performance at higher frequencies.To address this issue, another part was created wherein the tape was fully embedded (Figure 8). The transmission path was embedded, while the ground plane was surrounded by a lip of printed material in a channel to prevent ripples from forming. This part was designed with access windows to facilitate soldering. These windows were used to access the embedded copper tape for the purposes of soldering SMA connectors. The windows were then sealed with superglue before being manually filled with Vero White material and cured with a UV light for three minutes.Figure 8.

The windows were then sealed with superglue before being manually filled with Vero White material and cured with a UV light for three minutes.Figure 8. Fully-embedded micro strip transmission lineThis sample used dimensions identical to the previous micro strip, but had an impedance of less than 50 ohms due to being surrounded by dielectric (Figure 9). Additionally, this sample performed comparably well when compared to the micro strip line. Once again, the sample worked well up to ~4 GHz before it was subject to substantial losses.Figure 9. S-Parameter measurements for the fully embedded transmission line.4.4. Low-pass filterAfter successful construction of transmission lines, simple filter designs were explored. As described in Section 3.5, the samples were designed to have a -3 dB point at 1.6 GHz. The VW filter is shown in Figure 10.Figure 10.

As described in Section 3.5, the samples were designed to have a -3 dB point at 1.6 GHz. The VW filter is shown in Figure 10.Figure 10. Micro strip low-pass filterThe filter was simulated in Qucs producing the results shown in Figure 11, where the blue lineis S21 and the red line is S11. The Qucs simulation predicted a -3 dB point of 1.66 GHz.Figure 11. QUCs simulation of the micro strip low pass filter.The measured 3 dB point was 1.475 Ghz, which was within 200 MHz of the intended frequency and the simulation (Figure 12). Overall, the filter’s response is fairly similar to the QUCs simulation, suggesting that QUCs is a useful tool for simulating future micro strip filters. Part of the error in the calculation and simulation was due to error in an earlier estimation of the dielectric constant and loss tangent of VW.Figure 12. S-Parameter measurements of the low pass filterAdditionally, a model of the filter was produced using FEKO.

S-Parameter measurements of the low pass filterAdditionally, a model of the filter was produced using FEKO. The model and results of this simulation are presented in Figure 13 and Figure 14 respectively. This simulation used edge ports, as opposed to micro strip ports, to account for the finite ground plane dimensions, making it far more accurate. The -3 dB point predicted by the simulation, 1.66 GHz, matched the QUCs simulation. However, the FEKO simulation was more accurate than the QUCs simulation in predicting the shape of the response at higher frequencies. After obtaining a more accurate estimation of the dielectric constant and the loss tangent of Vero White, a new FEKO simulation of the edge port model was performed. After inputting these revised values, FEKO was able to predict the -3 dB point to within 15 MHz of the empirical results as seen in Figure 15.Figure 13. CAD model of the filter used for FEKO simulationsFigure 14.

CAD model of the filter used for FEKO simulationsFigure 14. S-Parameters produced from the FEKO simulationFigure 15. Zoomed view of FEKO results showing accuracy of the simulation valuesIn addition to creating a filter using VW, a filter made on Shore95 material was also fabricated (Figure 16). As with the VW sample, the filter was designed to have a -3 dB point at 1.6 GHz. Empirical results showed a -3 dB point at 1.565 GHz. Despite being the first attempt at a Shure95 structure, the -3 dB point was only 35 MHz from the design value, a 2.19% error. This sample illustrates the capability of AM technologies to produce useful RF structures from a variety of materials with known electrical properties.Figure 17. Tango Black low pass filter and corresponding S-Parameter results4.5. Edge coupled single resonator band pass filterAs described in Section 3.2, an edge-coupled micro strip band pass filter was designed to have a center frequency of 1.58 GHz.

Edge coupled single resonator band pass filterAs described in Section 3.2, an edge-coupled micro strip band pass filter was designed to have a center frequency of 1.58 GHz. After adjusting for the loss tangent of the VW material, a Qucs predicted a maximum S21 value of -5.21 dB at 1.58 GHz simulation (Figure 18). The predicted - 3 dB points were 1.47 GHz and 1.60 GHz.Figure 18. QUCs simulation of the band pass filter’s S-parametersThe empirical measurements showed a peak S21 of -5.23 dB at 1.535 GHz (Figure 19). The - 3 dB points were then 1.475 GHz and 1.595 GHz giving a 3 dB bandwidth of 120 MHz. The values closely match the Qucs simulation, and thus provide further of the loss tangent measurement.Figure 19. S-Parameter measurements of the band pass filter4.6. Micro strip patch antennasMicro strip patch antennas were fabricated on VW and TB substrates to function atapproximately 1.575 GHz in order to be able to receive GPS signals (Figure 20).Figure 20.

Rectangular Micro strip Patch Antennas on Vero White and Tango Black SubstratesThe FEKO simulation of the micro strip patch antenna utilized micro strip ports, which assume an infinite ground plane. The FEKO simulation of the power output from an excitation at the micro strip port predicted a maximum power output at 1.579 GHz (Figure 21).Figure 21. FEKO simulation of peak power for the patch antennaEmpirical S11 measurements using a network analyzer showed a local minimum at 1.585 GHz. The next closest data point was at 1.540 GHz. The maximum power output would be reflected by a minimum in the reflected power (S11). Thus, the empirical measurements validated the simulation results.The antennas were characterized using the anechoic chamber in the Virginia Tech Indoor Antenna Range. The antenna radiation patterns shown in Figure 22 Figure 23 were produced from measurements taken at 1.6 GHz on the Vero White patch antenna.

The antenna radiation patterns shown in Figure 22 Figure 23 were produced from measurements taken at 1.6 GHz on the Vero White patch antenna. The Tango Black patch antenna produced almost identical results and are therefore omitted. As can be seen in the figures, the measured antenna patterns closely mirror standard patch antenna radiation patterns.Figure 22. Co and Cross Polarized Far Field PatternsFigure 23. 3D representation of Co and Cross Polarized Far Field Patterns5. CLOSURE AND FUTURE WORKIn this paper, the authors demonstrate the feasibility to embed working RF structures into products fabricated via PolyJet additive manufacturing process.

CLOSURE AND FUTURE WORKIn this paper, the authors demonstrate the feasibility to embed working RF structures into products fabricated via PolyJet additive manufacturing process. Specifically, five RF structures (coaxial cable, micro strip transmission lines, low-pass filters, edge coupled resonator band pass filters, and micro strip patch antennas) created by embedding copper foil adhesive tape onto three PolyJet materials (Vero White Plus, Tango Black Plus, and Shore95, a composite material).Integral to the experimental technique was the characterization of the dielectric constant of these printed materials (Section 4.1).

In addition, simulation models were created using both Qucs and FEKO, and were used to verify the performance of the designed and fabricated RF structures, and thus validated their use as potential tools for the design of embedded structures.The embedded coaxial cable demonstrated very low loss well into the GHz range, and thus demonstrated the feasibility of embedding transmission lines. The fabricated micro strip line performed well up to about 5 GHz, and had an impedance of 55 ohms, which was close to the designed value of 50 ohms. A fully embedded micro strip had an impedance of less than 50 ohms than the open micro strip and worked well up to about 4 GHz, since it was surrounded by dielectric. The performance of the fabricated low-pass filters aligned well with the simulations; they were able to predict the -3 dB point to within 15 MHz of the empirical result for the Vero White Plus low-pass filter.

The predicted -3 dB point of the more flexible Shore95 filter was only 35 MHz from the design value, a 2.19% error. Thus, AM RF structures can be accurately constructed from a variety of materials as long as the dielectric constant and loss tangent are known. In addition, an edge coupled single resonator band pass filter operating from 1.475 GHz to 1.595 GHz and a micro strip patch antenna designed to function at 1.575 GHz were also produced.The ability to embed RF structures is critical for realizing an ideal AM system that enables the creation of multi-functional devices in a single process step. The authors look ahead to expand this work to take advantage of the multi-material capabilities of the PolyJet system. Specifically future designs may take advantage of 3D printing capabilities to print reconfigurable antennas or complex geometries with unique radiation patterns.6.

eating Implants from Allograft Bone using Subtractive Rapid PrototypingMatthew C. Frank, PhD, Ashish Joshi, MS, Shuangyan Lei, MS Department of Industrial and Manufacturing Systems Engineering Iowa State University, Ames, IA, USA Donald D. Anderson, PhD, Yuki Tochigi, MD, PhD, and Thomas D. Brown, PhD Department of Orthopaedics and Rehabilitation The University of Iowa, Iowa City, IA, USAAbstract This research involves the development of rapid manufacturing for bone implants using human allograft bone in a Subtractive Rapid Prototyping process. Using CT-derived CAD models of missing bone due to high energy trauma or tumor resection, surgical reconstruction could be improved with custom rapid implants made from natural bone. The bone “stock” material is of arbitrary shape and material distribution in the form of frozen donated cadaveric bones.

The bone “stock” material is of arbitrary shape and material distribution in the form of frozen donated cadaveric bones. Each is unique in shape and has highly anisotropic material properties; likewise for each final bone implant geometry and its material distribution. This work utilizes a PLY input file, instead of the more common STL, using color texture information that can be utilized for advanced process planning depending on whether the surface is fracture, periosteal or articular in origin. Moreover, we present a new PLY assembly model, called the Matryoshka model, where successively nested PLY files are used to designate changing material distributions in the donated bone; a method that could also aid in the use of multi-material additive RP systems.

Using color Matryoshka models and their subsequent color slice files; this work presents novel solution methods for the selection of implant harvesting sites and automated process planning for the physical rapid prototyping process. Early implementations using bone surrogate materials will be presented.IntroductionSegmental bone defects pose major challenges for orthopedic management. Whole sections of bone are oftentimes missing or must be surgically removed during the treatment of disease or trauma. In order for the bone as a whole to function properly, it is essential that the defect be filled with an implant that is both mechanically and biologically stable. Towards that end, shaping the implant from clinically relevant material can be challenging. Frequently, the surgeon sculpts these implants by hand to fit the defect, during surgery.

Towards that end, shaping the implant from clinically relevant material can be challenging. Frequently, the surgeon sculpts these implants by hand to fit the defect, during surgery. That task is prone to inaccuracies that can lead to complications, particularly for joint fractures, where a poorly filled bone defect can alter the joint mechanics, compromise implant fixation stability and ultimately cause the joint to degenerate. The field of rapid prototyping now offers several methods for creating implants from solid or porous materials, and from bioactive scaffolds.This paper presents the novel application of advanced Subtractive Rapid Prototyping (SRP) methods to the field of orthopedics [1,2]. The work was performed at the Iowa State University (ISU) Rapid Manufacturing and Prototyping Laboratory (RMPL), in collaboration with the Orthopedic Biomechanics Laboratory at the University of Iowa (UI).

The objective of this research was to develop algorithms that would facilitate machining of patient-specific bone implants from frozen donor cadaveric bones (allografts) as the stock material. Implant made from allografts could have a unique density distribution throughout its volume. Having a density distribution on the implant similar to the one at fracture site could allow the parent bone to integrate with the implant effectively compared with other bio-implants. Also, in order to facilitate effective healing, it is necessary for the implant to maintain its initial fixation stability. This could be achieved by providing surface specific characteristics (textures) on the machined implants.

This could be achieved by providing surface specific characteristics (textures) on the machined implants. Hence the algorithms designed in this research aid in choosing an appropriate implant location within the donor bone to impart unique density distribution and choosing process parameters that facilitate implant machining with customized surfaces (textures) to increase the implant’s fixation326stability. The CAD geometries of segmental defects (residual voids/implants) were extracted and designed using fracture reconstruction planning software developed at the University of Iowa (FxRedux) [3,4] to analyze patient CT data.

The corresponding custom defect filling implants were fabricated using an SRP technique called CNC-RP at Iowa State University.The clinical need: custom-fit allograft bone for implantationEven in the current era of sophisticated bone grafting procedures, advanced synthetic biomaterials and bioactive/tissue engineered implants, refined capabilities for restoring soft tissue coverage, and highly evolved distraction osteogenesis techniques [5], treating segmental bone defects presents a major challenge. To date, most attention in this area has focused on mid-shaft long bone defects, where the principal reconstructive objective is to achieve bone healing with nominal preservation of limb length and alignment. While shape matching between the graft and the recipient site is always desirable in principle, many mid-shaft fractures are relatively forgiving in that regard. Various other bone defects, by contrast, place a much higher premium on close geometrical matching of the graft.

Various other bone defects, by contrast, place a much higher premium on close geometrical matching of the graft. For example, bone defects associated with severe articular or peri-articular fractures (i.e., fractures near a joint such as the knee or hip) require a substantially higher degree of reconstruction accuracy than is the case for the mid-shaft defects, owing to the need for stable, congruous articulation of the joint surface. Bone healing of an articular fracture in other than closely anatomic position predisposes the joint to secondary arthritis, a major contributing factor to poor outcomes, whose morbidity frequently approaches that of amputation [6].At the local macroscopic level, all fractures possess individual geometric signatures.

Current synthetic implant or grafting strategies for achieving healing of segmental defects offer only limited opportunity to address individualized defect geometry, since they have evolved mainly for situations (mid-shaft defects) where close reconstruction of local geometry is not particularly critical. Using conventional methods, there has to be primary reliance on fixation hardware to hold the respective bone surfaces in the desired nominal apposition, with the implant or graft making at best local spot contact with the recipient bone, and with appreciable gaps existing across much if not most of the intended-union interfaces. Even with the most advanced contemporary fixation trained orthopaedic peri-articular fractures (especially in the presence of segmental defects) pose a severe biomechanical challenge, that often is not well resisted by usage of conventional bone grafts [7].

Virtually all contemporary synthetic implant materials, all tissue engineered defect-filling constructs, and especially all variants of bone grafts would have better prospects for achieving optimal outcome, if they began from a condition of closely fitting the local geometry of the recipient bone surface(s).(a)Rotary IndexerEnd MillRoundAxis of RotationOpposing 3-Jaw chucksthe hands of highlyincomminutedtraumatologists,(b)(Side View) (Side View)(5)(1)(6)(2)(3)Rapid manufacturing using CNC-RP(4)Final Steel PartCNC-RP is a fully functional Subtractive Rapid Prototyping (SRP) process using a standard 3-axis CNC milling machine with a 4th axis for multiple setup orientations.

The CNC-RP approach features completely automated fixture planning, tooling and setup planning, including generation of NC codeFigure 1 - (a) CNC-RP setup; (b) iterative steps of layer based toolpath planning from multiple angles, using sacrificial supports to fixture the part to the remaining stock327for creating a part directly from a CAD file [1,8,9]. The use of a rotation axis eliminates the need for re- clamping of the part, a common task in conventional fixturing methods [10,11]. For each orientation, all the visible surfaces are machined while a set of sacrificial supports keeps the part connected to the uncut end of the stock material. Once all the operations are complete, the supports are severed (sawed or milled) in a final series of operations, and the part is removed [12]. The setup and steps to this process are illustrated in Figure 1.

The setup and steps to this process are illustrated in Figure 1. The manufacturing of biomedical implants provides a challenge very well suited for CNC-RP, especially due to the fixturing issues and the need for specialty materials (in particular, human allograft bone). Figure 2 shows a fragment from a human tibia which was reverse-engineered from a CT scan and then rapid-machined from clinically relevant materials using the CNC-RP process.Bone Implant HarvestingThe authors have been developing methods to harvest custom bone implants from donor bone, with implant geometries derived from CT scanning of patients’ bones [2]. In order to effectively create patient-specific implants from a donor bone, one needs to find the optimal location within the donor bone from which to harvest the implant. A challenge in this approach is that donor bone is unlike a homogeneous artificial bone material, where implant could be harvested from anywhere within the stock material volume.

A challenge in this approach is that donor bone is unlike a homogeneous artificial bone material, where implant could be harvested from anywhere within the stock material volume. Natural bone has heterogeneous internal structure, and the site of implant harvest needs to respect the realities of internal bony density distribution; not all harvest locations that are geometrically feasible will be clinically suitable. Figure 3(a) shows a cross-sectional view of a femur bone, showing the spongy, low density trabecular bone in the middle, versus the high-density cortical bone on the outside. A surgical bone implant might need to have a bone material density that is highly, or almost entirely cortical bone, or it may need a particular distribution of densities, or it may simply need to try to match the general gradient directions.

Regardless, there needs to be an efficient way to represent the complex material properties given piece of such “stock material”.theFigure 2 - Implant prototyping from CT scan to machined porous metal (Trabecular MetalTM ) and native bone (bovine)(a)(b)(c)Figure 3 – Comparison of bone to stacking doll models, (a) A cross-sectional view of a femur bone, Matryoshka shell model; (b) Set of Matryoshka dolls, in order of size and (c) Cut-away view of nested Matryoshka dollsTo meet this challenge, we propose the idea of a Matryoshka shell model as a way to describe the bone density distribution, using discrete regions generated from CT image data. As a simple example, Figure 3(b) shows a set of Matryoshka nested dolls. The salient characteristic of Russian nesting dolls is that the size of each nested doll decreases in order to place one inside the other, as shown in Figure 3(c).

The salient characteristic of Russian nesting dolls is that the size of each nested doll decreases in order to place one inside the other, as shown in Figure 3(c). Although models generated from bones will not exactly follow the monotonically decreasing regions of these dolls, the general concept of nested shells is the underlying principle.328The Matryoshka Shell Model The task of defining density function is not trivial, since bones have a complex heterogeneous structure. There have been numerous research efforts to model the density of bone. The Hounsfield Unit (HU), which quantifies the X-Ray attenuation, indicates the varying bone density levels; High HU corresponds to high density, and vice versa. HUs are also associated with grayscale of the CT slice image, which carries intensity information.

HUs are also associated with grayscale of the CT slice image, which carries intensity information. By setting threshold HU values, pixel values below given threshold values can be set to be pixels-of-interest, while HU values above the threshold values can be set to be background pixels. Figure 4 shows an example of distinguishing pixels of interest from a CT slice image, using the HU threshold method divided into 5 different regions bounded by different contours. When each increasing contour shell is created, it is assumed that the pixels within that region have a common singular HU around the thresholding value. Hence, the continuous bone density function exhibited on the CT slice can be discretized into a step function. The 3D Matyroshka model created by iterative thresholding operation on a series of CT slices and stacking them together is saved as a single PLY file. This PLY file can be used for determining the location in the donor bone from which the implant can be machined.

Although the space can appear to have infinite solutions for positioning and orienting the implant, we wish to reduce the degrees of freedom for searching based on known characteristics of human long bones. Usually human long bones (femur, tibia,329humerus, etc) are attractive as donor bones. If shape similarity is considered between a long bone and a cylinder, a long bone generally exhibits increasing density away from the axis radially, or the medullary canal. This gives a basis for defining the “center line” of the bone. Before initiating a harvesting search within the donor bone, it is first important to align the “center lines” or axes of the patient bone and donor bone. Thus, the density distributions in the implant and the donor bone would generally align radially along the same axis. When one traverses a set of cylindrical coordinates in the donor or patient bone space, the gradient directions will generally align (i.e.

When one traverses a set of cylindrical coordinates in the donor or patient bone space, the gradient directions will generally align (i.e. moving away radially will increase density in both, and vice versa). Hence, our simplified harvesting search involves implant rotation about and translation around the center line in the donor bone space.Creating a discretized slice modelBefore the iterative search begins, the Matryoshka model is sliced about the center line. The colored slice can be considered as a boundary; the region within each color boundary represents its density and is given one specific value. In this manner, all elements contained between two adjacent shells are set to the same value, other than the first shell which contains the medullary canal and this is not a feasible region for bone harvesting. As shown in Figure 6(b), values a, b, c, and d represent the bone densities from different regions in the slice model.

As shown in Figure 6(b), values a, b, c, and d represent the bone densities from different regions in the slice model. Next, a spider cell structure is used to discretize each slice into a grid of sectors about the Z-axis, with an interval angle α, and with each sector further divided by grid elements with interval h, as shown in Figure 6(c). Each grid element is assigned a specific value indicating the density of the region encompassed by the corresponding grid element.

Each grid element is assigned a specific value indicating the density of the region encompassed by the corresponding grid element. In general, smaller α and h will result in a more accurate and finer discretization structure, but at the expense of increased computation time (The grid spacing for both α and h are shown excessively coarse in Figure 6(c) for clarity; in practice, they are 1-5˚ and 1-3mm, respectively).Figure 6 - (a) One cross-section from a Matryoshka shell model; (b) Color mapping for different density regions; (c) Discretized density regions.For each slice, an array is used to indicate the presence of the shell in the grid structure. If any side of the grid element intersects the chain, a specific value is assigned to represent the existence of a specific shell in that element at a specified density, with values of a, b, c, d used to represent the existence of Shell 2, Shell 3, Shell 4 and outer surface Shell 5, respectively.

If no shell intersects any grid element, a value of 0 is assigned to the corresponding element in 2D array as shown in Figure 7(b).Since all grid elements that fall within the surface boundaries of the bone will be assigned to one of the specific density regions, those grid elements temporarily indexed in the array with value of 0 need to be modified to represent the density correctly. For example, after filling in the 0 elements of the array, the updated row of the array is as shown in Figure 7(c).330Figure 7 - Section chains intersecting the grid elements, (a) One spider cell region, (b) Shell boundary detection, and (c) Element fillDensity score and similarity score calculationThe goal is to assign a quantitative score of “goodness” for a candidate location for bone implant harvesting.

This will be focused on the density of the implant, regardless of whether the preference is simply for high overall density, or for a highly similar distribution of densities between donor and patient. To begin, after discretizing the slice model into the grid structure and assigning a density value in each element, the area of each element is determined based on parameters α and h. To calculate the overall effectiveness score for the entire provisional implant position, density metrics are calculated for each chain of each slice for the entire implant, as follows:Let N be the total number of slices for the implant, for i slices from 0 to N-1. Let j , ranging from 0 to 3, represent the four bone density regions. Then, let Sij indicate the area of the different density regions j on slice i. Recall that the density regions 0, 1, 2, 3 represent the range from low-density cancellous bone region to the highest density cortical bone region.

Whereas Density is an aggregate score for the entire implant, Similarity is evaluated slice by slice. Hence, although one could achieve a high overall density score by having some portion of the implant gain density at the expense of another portion losing density; the Similarity score will be affected more locally.Conceptually, the simplest approach is to conduct an exhaustive search of the entire donor bone space, in order to determine the optimal location for the harvested implant. This exhaustive search involves rotating the implant about the Z-axis, and translating the implant up and down in the Z-axis direction, while moving the implant near to or far away from the Z-axis radially. In other words, the implant is moved at331the granularity of the spider grid structure, throughout the entire donor bone space. For each iteration, the Density and Similarity are calculated, and then both values are normalized.

Using CNC-RP, we currently have solutions for automatically machining a part from simple stock, with homogenous properties. The problem now is that we are faced with a “part”, the implant, where its functionality, both density and surface characteristics can have significant impact on biocompatibility and performance. To this end, the following section illustrates how one can find surface specific machining operations based on a colored polygonal model; allowing us to deliver custom surfacing on the implants without human intervention in process planning.Process planning for calculating sub-surface specific orientationsAfter determining the location for harvesting the implant, setup orientations can be calculated for machining the implant with customized surfaces. In previous work for CNC-RP process planning, it was only deemed necessary that the entire surface of the part model be machined after all setup orientations were completed.

In previous work for CNC-RP process planning, it was only deemed necessary that the entire surface of the part model be machined after all setup orientations were completed. This problem of calculating the set of setup orientations for machining the entire part is classified as a Set Cover problem, where the whole surface of the part model visible in the range of 0° to 360° is included in the universal set (Figure 8). The algorithms designed for the CNC-RP ensure that each portion of the part surface visible in the range of 0°Figure 8 - Setup orientations using STLto 360° is machined from at least one setup orientation from the solution set. Referring back to the CNC-RP introduction in Figure 1, this basically entails determining the setups for the machining process about the rotary axis, or number of times we need to stop and deploy a set of toolpaths.

Due to the lack of sub-surface identification on the STL file, the previous algorithms for calculating setup orientations were designed to target the entire model geometry, but do not create different finishes on each sub-surface. Instead, if one could pass information about different surfaces on the bony anatomy of the implant then customized machining operations could be implemented for each. In this work, color PLY files are used to designate surfaces as Fractured, Articular, or Periosteal in origin, passing this color of the facets onto the slice polygons (Figure 9). The basic Set Cover approach is used here, but with the significant difference of achieving set cover for each sub-surface individually, rather than for the entire model. Thus, in order to target individual sub-surfaces, setup orientations have to be chosen such that they are aimed at sub-surfaces individually (Figure 10) rather than at multiple sub- surfaces together.

In this work, setup orientations specific to the articular/periosteal/fracture sub-surfaces are designated with subscript (cid:2242)a/p/f. The process planning algorithms developed for choosing sub-surface- specific setup orientations consider the primary variables of 1) Surface Visibility (SV), 2) Surface Reachability (SR), and 3) Normal Deviation (ND) (Deviation of a candidate setup orientation from the average sub-surface normal).Figure 9 – Color PLY file for implant geometry332Sub-surface Visibility (SV)Visibility is a necessary condition for a given sub-surface to be machined by CNC-RP. For example, if the objective is to machine the articular sub-surface, a setup orientation from which articular sub-surface is visible must be chosen for machining it.

For example, if the objective is to machine the articular sub-surface, a setup orientation from which articular sub-surface is visible must be chosen for machining it. Hence the primary aim would always be to maximize the visibility of the sub-surface for which setup orientation(s) is being chosen.Sub-surface Reachability (SR) In order to ensure that entire part can be machined, it is necessary that every visible sub-surface on the part also be reachable from at least one orientation. For example, there could be an instance where a certain percentage of a sub- surface is visible from a given orientation, but is not reachable because of inadequate tool length. Thus, the total reachable perimeter is calculated by comparing the maximum available tool length against the perpendicular distance from each visible point to the tangent line at the given orientation (Figure 11).

Hence, the aim here is to maximize the Surface Reachability while choosing setup orientations.Figure 10 - Setup orientations targeting individual sub-surfacesNormal Deviation(cid:4666)ND)One of the factors to be considered is the texture directionality created on a sub-surface. For example, in order to have stable implant fixation, it is desirable that we can impart a textured surface onto it to allow better bite into the existing site. This ability increases as we can align toolpaths with a more normal apposition to the desired surfaces. Hence, minimizing the Normal Deviation (ND) is sought (Figure 12). This is not only true when attempting a rough surface, but also for a smooth surface, where an increase in ND generally increases the scallop height.

This is not only true when attempting a rough surface, but also for a smooth surface, where an increase in ND generally increases the scallop height. Ironically, the same is needed whether an intentionally rough or smooth surface texture; hence, the aim here is to have minimal Normal Deviation (cid:4666)ND(cid:4667) for setup orientations.Figure 11 - Surface Reachability Tool Length L < Depth D,Goodness Measure for a Setup Orientation The previously defined measures can each aid in defining a good or bad setup orientation from all candidates; however, we desire an overall composite measure to drive toward an optimized set of setups. For a given candidate setup orientation, we propose a quantitative goodness measure as defined by four variables, namely; 1) Toolpath Containment (TCO); 2) Toolpath Crossover (TCR); 3) Toolpath Redundancy (TR), and 4) Normal Deviation.

While Normal Deviation simply is the difference between the candidate orientation and the average sub-surface normal, the other three variables are functions of Surface (SR) defined Visibility(SV) and Surface ReachabilityFigure 12: Normal Deviation between a candidate setup orientation and average surface normal333previously. The following presents each proposed variable used in the goodness calculations.Toolpath Containment (TCO)Toolpath Containment (TCO) is the idea of machining the visible and reachable sub-surface of primary interest. Since the overall goal is to choose the minimum number of setup orientations to machine a sub- surface, a setup orientation with maximum percentage TCO will always be chosen to machine the maximum sub-surface.

For example, TCR from the periosteal to fracture sub-surface means that both these sub-surfaces are machined when the intention was to machine only the periosteal sub-surface. TCR can have harmful effects on the implants functionality. For example, TCR from the periosteal to the fracture sub-surfaces can reduce the primary fixation stability of the implant. TCR from fracture to periosteal sub-surface would create a rough texture on the periosteal sub-surface, which could irritate the overlying soft tissues. Regardless, the goal is simple; machine each sub-surface with customized toolpaths, and avoid machining other sub-surfaces while doing so.

For example, Redundant machining on periosteal or articular sub-surface would just be inefficient, since additional smoothing of those sub- surfaces is insignificant. However the machining time could impact the practical use of this technology in a production setting (cost, machine capacity, etc.). Furthermore, redundant machining of fracture sub- surface could also reduce texturing effects (or ablate them completely). This would potentially lead to reduction in fixation stability of the implant.

Additionally the customized sub-surface roughness helps to increase the implant’s initial fixation stability and maintains its biomechanical compatibility. It is hypothesized that this will reduce healing time in the short term and reduce the likelihood of longer term issues like PTOA. The algorithms for determining the implant harvesting site and sub-surface specific setup orientations were implemented in computer software using C++ and are presented as follows.Harvesting software implementationThe algorithms developed in this work have been implemented in C++ and are graphically displayed using OpenGL. To illustrate the implementation of these analyses, two sets of human tibia bone CT scan slices are used. These slices were first imported into ITKsnap (open source software) and saved as a voxel array DICOM file. Next, the DICOM data were loaded into Matlab, and the five Matyroshka shells were generated.

Next, the DICOM data were loaded into Matlab, and the five Matyroshka shells were generated. The created shells were then saved into a .ply file in MATLAB, and imported to RapidForm software for post-processing. Post-processing included correcting geometric errors in the .ply file (holes, spikes, etc.).

(a) Initial location of the implant within the donor bone; (b) The location of the implant is updated by rotating about the Z-axis by 103.5(cid:2925), translating along the Z-axis direction by -4.0 mm, and moving radially by 1.0 mm(a) (b) (c)337(d) (e) (f) (g)Figure 19 - An example of implant harvesting using CNC-RP. (a) Harvested implant within the bone; (b) The potted bone is clamped in CNC milling machine; (c) Section cut of the bone; (d) Bone section with sacrificial supports, (e) Rough machining; (f) Finish milling; (g) Finished implantProcess PlanningThe above-described algorithms for calculating sub-surface specific setup orientations were implemented in C++ and an OpenGL user interface, and were tested on an Intel Core2Duo, 2.8 GHz PC, and running Windows 7. The software accepts colored 2D slice files from 3D PLY models as input and output setup orientations necessary to machine the implants with customized surface roughness.

Compared with the traditional way of hand-creating an assembly model, this method could potentially be completely automated, given a set of parameters. This work illustrates how the Matryoshka model can be used to plan harvesting locations for creating custom bone implants from within actual human donor bones, and it develops an approach to calculate a Density score and Similarity score for an arbitrary provisional implant harvest site, to evaluate the overall effectiveness of that harvest site. Future work with the Matryoshka approach could be to develop a better harvesting solution for irregular and/or flat bones with no proper axis definition.

Future work with the Matryoshka approach could be to develop a better harvesting solution for irregular and/or flat bones with no proper axis definition. This method also could be used for industrial components which may have less amorphous shapes, and which could be printed using existing additive systems such as LENS or polyjet printing for objects with gradient material properties, for example.Process Planning Additionally this paper also introduces a new method for calculating setup orientations to create implants with customized surfaces. Texturing the fracture sub-surface could reduce implant/host bone interfacial movement and increase stability, while a smooth periosteal and articular sub-surfaces could enhance biomechanical compatibility. This technique for calculating setup orientation has the potential to produce implants with improved performance, and consequently to improve patient outcomes.

This technique for calculating setup orientation has the potential to produce implants with improved performance, and consequently to improve patient outcomes. The proposed algorithms could also be modified for industrial purposes such as in applications where the number and/or types of sub-surfaces present on the part may be more numerous, for example, when a variety of GD&T callouts are defined.Acknowledgements This work was supported by grants from the NIH/NIAMS (P50AR055533 and R21AR054015), the Roy J. Carver Charitable Trust at the University of Iowa, State of Iowa Economic Development appropriations to the Board of Regents under the Grow Iowa Values Fund, and the Musculoskeletal Transplant Foundation. The authors would also like to recognize the technical support of Mr. Gary Ohrt and Dr. Thad Thomas at the University of Iowa.[1] Frank, M.C., Wysk, R.A., and Joshi, S.B.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperCREATING TOOLPATHS WITHOUT STARTS AND STOPS FOR EXTRUSION- BASED SYSTEMSAlex Roschli*, Brian Post*, Phillip Chesser*, Michael Borish*, Lonnie Love*, Seokpum Kim*Oak Ridge National Laboratory, Manufacturing Demonstration Facility, 2350 Cherahala Blvd, Knoxville, TN 37932AbstractToolpath generation for extrusion-based additive manufacturing systems, called slicing, involves operations on polygonal contours that are derived from an STL file. Slicing generates multiple paths per layer (both closed-loop and open-loop) that are designed to optimally fill the space outlined by the polygon(s). In the course of printing a layer, the extruder must start and stop, the tip must be wiped, and the extruder must travel between paths without printing.

In the course of printing a layer, the extruder must start and stop, the tip must be wiped, and the extruder must travel between paths without printing. Any amount of time the printer spends moving without printing is considered wasted time because the part isn’t being constructed. In addition, the start/stop point, known as the seam, is often a blemish on the surface of the part that contributes to weaker material properties. Therefore, a single path for creating multi-bead walled structures is desirable because it would save machine time and create parts with better surface finish. This paper will cover one method of modifying the CAD file and slicing engine to allow for parts to be printed without starting and stopping the extruder.IntroductionFused Deposition Modeling (FDM) is an extrusion-based 3D printing process [1] that can be used to additively manufacture objects by extruding metal [2], polymer [3], and concrete [4].

Extrusion-based systems use toolpaths to direct the deposition head to deposit material, layer by layer, in desired locations. When printing simple geometries, the process is straightforward. The tip of the printer head has to be wiped in between layers to clear excess material, but otherwise, printing is continuous. However, when more complex geometries are printed, it is not uncommon for toolpaths to contain multiple starts and stops. For example, any holes or gaps within the printed part are often achieved by printing the contour of the hole, so that the area surrounding the hole has a smooth surface finish. The contour of the hole is usually separate from the outer edge of the part. Therefore, the printer head has to stop printing, move to the edge of the hole, and start printing again. The higher the number of distinct edges and contours a geometry contains, the more starts and stops are introduced into the toolpath.

The higher the number of distinct edges and contours a geometry contains, the more starts and stops are introduced into the toolpath. These starts and stops necessarily increase print time, as all the time spent not printing is time spent not producing the part. Therefore, to increase the efficiency and appeal of FDM, it is necessary to develop a way to continuously print complex geometries while preserving surface finish with minimal starts and stops.The standard method for generating extrusion-based system toolpaths is a purely geometric-based approach called “slicing” [5,6]. The name slicing comes from the process of slicing the file to be printed into layers. The part, which originates in CAD (computer aided design) software (such as Solidworks or Autodesk Fusion 360), is exported from the CAD software as an STL (stereolithography) file. An STL file represents the surface geometry of the CAD file using triangles to approximate all curves and features [7].

An STL file represents the surface geometry of the CAD file using triangles to approximate all curves and features [7]. This STL is loaded into the slicing software to be sliced into layers.1113Slicing works by intersecting a horizontal plane (in plane with the build surface) with the STL file and then incrementing the height by the desired printing layer height. Each slice becomes a layer represented as a polygon or polygons to show the boundaries. The next step of slicing is to fit toolpaths to the layer to optimally fill the layer. Toolpaths are generated specific to the user- defined settings in the slicer such as bead width, infill density, and number of perimeters. Closed- loop paths are generated first by polygon offsets of the layer’s bounding polygon. Once all closed loop paths are generated, usually called perimeters and insets, skins and infill are used to fill the remaining void.

Once all closed loop paths are generated, usually called perimeters and insets, skins and infill are used to fill the remaining void. These can be generated in many different ways, but the simplest approach is to apply a grid across the region to be filled and apply toolpaths to the grid.The final step of slicing is to output the toolpaths as G-Code, which is code that represents the instructions for the machine about how to print the desired part. G-Code has some standardizations for common codes but is largely machine-specific. Each machine has its own unique hardware i/o that can be enabled and disabled as needed during the print. The slicing software must know about each system’s capabilities and format the G-Code output accordingly.Starts and StopsStarts and stops occur when one toolpath ends and the extruder stops extruding to move to the next toolpath.

The time spent moving between the end of one path and the start of the next, called traveling, slows down the print and therefore necessarily wastes time. Travel time is considered wasted time for manufacturers because the machine is not adding material to grow the part. While some part geometries and some manufacturing systems require travel time to ensure the part is successfully printed, minimizing unnecessary travel time is always ideal [citation?]. Therefore, the challenge for any system is to develop approaches that achieve successful prints while minimizing starts and stops.The start and stop point, known as the seam (Figure 1), creates a non-uniform region on the surface of the part where the extruder must stop and the tip must be wiped off to clear excess material from the nozzle. This seam is not only rough, but it creates a weak “joint” in the part [8]. At the start/stop point, the bead of polymer is disconnected.

This seam is not only rough, but it creates a weak “joint” in the part [8]. At the start/stop point, the bead of polymer is disconnected. It’s a butt joint where fibers don’t flow across the joint, causing a significant loss of strength.Figure 1: Starts and stops, a weak point on Big Area Additive Manufacturing (BAAM) prints1114The start of a path for a pellet-based extruder requires a pause, or a slow movement, to build up pressure in the extruder so that the initial extrusion will bond to the substrate or previous layer. The end of the path has the opposite problem. The release of pressure as the extruder slows down causes material to drool out of the extruder nozzle. This can be corrected by moving slowly for the last few moves of the path or wiping the tip across the part, both of which take time.The Single PathCurrently, single path printing exists through a method called spiralizing [9].

Spiralizing works by removing the independent closed loop paths of a layer to make just one continuous closed loop path for each layer. This is done by taking the layer polygon and offsetting the outer boundary inward by one bead width. The result is a single closed loop path to represent the external geometry of the layer. Figure 2 shows a circle and the resultant toolpath when spiralizing.Figure 2: A circle and the spiralize toolpath that it generatesIf every layer of a spiralize print is planned so that the start and stop is at the same location, then the start and stop can be removed by continuing the print from one layer to the next. This is done by slowly incrementing the Z height throughout the layer so that after one full layer has been traversed, the Z height has changed by one-layer height.

This is done by slowly incrementing the Z height throughout the layer so that after one full layer has been traversed, the Z height has changed by one-layer height. Figure 3 below shows a chair that is printed with spiralize.1115Figure 3: A chair printed without starts and stops, called spiralizingSpiralizing doesn’t work for dense structures because there can be no starts and stops, which means no infill or solid regions and no two-bead wall objects. It also means printing just one object at a time. Because of these constraints, spiralizing is often called spiral vasing because it’s used to print vase-like objects that just need the outer contour without internal spars or support structures.The spiralize approach can be taken a step farther to encompass a full layer, if the extruder has a path to traverse the layer without crossing itself or starting and stopping. Take for example a simple square with the center removed, Figure 4.

Take for example a simple square with the center removed, Figure 4. The standard tool pathing approach for this results in the toolpaths seen in Figure 5. This object has three distinct paths that require starting and stopping the extruder and traveling. If the extruder got to a corner and turned around and traveled back on itself, it could return to the start point while tracing all paths and never starting/stopping (Figure 6).1116Figure 4: A square with hollow centerFigure 5: Standard tool pathing for a hollow square with two bead wall1117Figure 6: Square with tool pathing forced to double back to create a single pathThis same approach can be applied to larger structures, like trusses, that can be constructed with multiple closed loop paths. The number of times the extruder must turn around on itself is the number of closed loop paths minus one. If the part is designed such that every layer starts and stops at the same location, a spiralize approach can be implemented.

If the part is designed such that every layer starts and stops at the same location, a spiralize approach can be implemented. Figure 7 below shows a figure eight with two cuts to reduce the print from three paths to one path.Figure 7: A figure eight printed as one single path1118Single Path STLThe simplest method of creating a single path is to simply design the object to be printed as a single path. This can be done by inserting small cuts along the wall sections to break up the closed loop paths and force the extruder to double back on itself (Figure 8). These cuts are very small, just 0.02inches (0.5mm) wide, which is 1/15th the size of the nozzle diameter. A cut this small prevents the trusses in the structure from being connected in the tool pathing (when optimization is disabled) but doesn’t leave a gap during the actual print.

A cut this small prevents the trusses in the structure from being connected in the tool pathing (when optimization is disabled) but doesn’t leave a gap during the actual print. This forces the extruder to double back on itself and create one large closed loop instead of individual paths.Figure 8: An object with small cuts to force the travel in a single pathSlicer UnoptimizationIn order for printing to be as fast as possible, tool paths are automatically optimized by the slicer to save time and eliminate unnecessary moves. For example, paths of a length smaller than the nozzle diameter can’t be printed and should be eliminated. The same goes for holes that are smaller than the nozzle diameter; they get eliminated because the holes add more time to the print than they remove [10].This removal of small holes creates a problem for the small slices that were added to the STL to create a single path.

The slices are removed by default during the optimization step because the slicer thinks it’s saving time and optimizing out something that won’t work. Figure 9 shows what happens when the STL from Figure 8 is sliced with the optimization enabled. The small cuts get ignored, or optimized out, and the layer is printed normally by breaking up all the sections into small closed loop paths. To fix this, a setting was added to the ORNL Slicer that allows the user1119to skip the optimization step during inset generation. Figure 10 shows the resultant G-Code generated from the STL in Figure 8.Figure 9: The result of slicing software optimizing out all the small cuts.

Figure 10 shows the resultant G-Code generated from the STL in Figure 8.Figure 9: The result of slicing software optimizing out all the small cuts. The pink areas are starts/stops and the light blue lines are travel moves.1120Figure 10: The resultant toolpaths when optimization is disabledResultsTo compare the two approaches, printing the traditional way with starts and stops and the single path approach, the object seen in Figure 8, a representative boat hull, was printed both ways on a Cincinnati BAAM system. The object is 60” long and 42” wide. The objects were made with a 0.3” nozzle and a 3” long tip wipe. Figure 11 below shows the print with traditional pathing including starts and stops. Figure 12 is a closeup of the starts and stops to show the roughness of each seam. Figure 13 shows the single path print and Figure 14 shows a closeup of the cuts where the extruder doubles back on itself.

Figure 13 shows the single path print and Figure 14 shows a closeup of the cuts where the extruder doubles back on itself. The cuts are small, just 0.02” wide, but Figure 14 shows that the beads still have a small amount of overlap preventing an actual gap from occurring during printing.1121Figure 11: Print with traditional pathingFigure 12: Closeup of the starts and stops in Figure 111122Figure 13: Print with a single path, no starts and stopsFigure 14: Closeup of the inserted cuts to create the single pathTable 1 shows a data comparison of the two prints. The travel distance for the traditional pathing is the XY distance between the start and stop of each move including the transition to the next layer.

The travel distance for the traditional pathing is the XY distance between the start and stop of each move including the transition to the next layer. The single path only has one travel move, from the end of the tip wipe in a path to the start point of the next layer.1123Table 1: Comparison of printing with traditional pathing using starts and stops and a new approach using a single pathNumber of Printing Paths Number of Travel Moves Travel Distance (in) Layer Time (s)Traditional Pathing 16 16 158.88 129.36Single Path 1 11 3 109.57ConclusionAn approach for creating a single path to traverse a layer of a 3D printed part was presented. This approach is meant to save time during the printing process by eliminating starts and stops and non-printing moves. There are geometric limitations that prevent this method from being applied to all 3D printed objects, but the methodology is very applicable to large format processes such as BAAM.

There are geometric limitations that prevent this method from being applied to all 3D printed objects, but the methodology is very applicable to large format processes such as BAAM. For the test print of a boat hull, the single path approach saved more than 15% (19.79 seconds per layer) of the time the traditional path took. This time savings is geometry-dependent, so not all prints will see 15%-time savings. Based on the results of this study, it can be inferred that objects with a lot of individual short paths will see a large time savings and parts with less paths will see less time savings. Further studies investigating the use of this method on a wider variety of part geometries are required to conclusively demonstrate the proposed time savings.

Further studies investigating the use of this method on a wider variety of part geometries are required to conclusively demonstrate the proposed time savings. Additionally, future work will entail mechanical testing, to see if there are differences in material strength, and software algorithm development to remove the manual slicing in CAD.AcknowledgementsThis material is based upon work supported by the U.S. Department of Energy, Office ofEnergy Efficiency and Renewable Energy, Office of Advanced Manufacturing, under contract number DE-AC05-00OR22725.References[1] Comb, J., et al. (1994). FDM Technology Process Improvements. Solid FreeformFabrication, 5, 42-49[2] Mireles, J., et al. (2012). Fused Deposition Modeling of Metals. Solid Freeform Fabrication,23, 836-845[3] Duty, C., et al. (2017), "Structure and mechanical behavior of Big Area AdditiveManufacturing (BAAM) materials", Rapid Prototyping Journal, Vol. 23 No. 1, 181-189.[4] Chesser, P., et al. (2018).

EEP BEHAVIOUR OF POLYAMIDE IN SELECTIVE LASER SINTERING Eef Moeskops, Nico Kamperman, Bart van de Vorst, Rik KnoppersTNO Industrial Technology De Rondom 1, P.O. Box 6235, 5600 HE Eindhoven, The NetherlandsReviewed, accepted August 19, 2004ABSTRACTThe use of Rapid Prototyping technology to produce fully-functional end use products (Rapid Manufacturing) requires understanding of the material's time dependency. Creep is the time dependent strain as a result of a constant load. Creep behaviour of sintered polyamide-12 processed by the Selective Laser Sintering process has been investigated. To study the effect of temperature on this material, tests have been performed at room temperature as well as at elevated temperatures of 60°C and 100°C. Although the SLS processed material is more sensitive to creep than most other thermoplastics it is shown to be affected less than injection molded polyamide-12, which has been tested as a reference.

Although the SLS processed material is more sensitive to creep than most other thermoplastics it is shown to be affected less than injection molded polyamide-12, which has been tested as a reference. The results show an almost linear behaviour on a logarithmic time scale, which enables the creep behaviour of SLS processed polyamide-12 to be predicted for at least 11 years.INTRODUCTIONRapid Manufacturing (RM) is the production of fully-functional end-use products using Rapid Prototyping technology. One of the techniques that was found to have potential to be utilised as an RM process is Selective Laser Sintering (SLS). With SLS, a powder - having particle sizes in the range of 20~100 µm - is spread out onto a build platform that steps down for each layer of the model. A CO2 laser selectively heats up the powder to a temperature just above the melting point, bonding the particles in one layer together as well as to the layer below.

A CO2 laser selectively heats up the powder to a temperature just above the melting point, bonding the particles in one layer together as well as to the layer below. The whole process is kept inside a heated chamber at a temperature close to the melt temperature of the powder, so that the laser needs to impart only a small temperature increase to ensure bonding. The layered manufacturing and the polyamide material are at the base of some characteristic RM-material properties, such as: brittle failure, anisotropy, temperature dependency and moisture absorption. These material properties can be analysed using 'short-term' tests such as tensile tests. Designing for RM also requires understanding of the time-dependent 'long-term' properties. After all, the product needs to meet the requirements even after a period of time in service.An important 'long-term' property is tensile creep, which is the strain (∆l+ ∆c) as functionof time (t), as a result of a constant load (F).

See Figure 1.60l0l0+∆ll0+∆l+∆c∆l∆lF∆cFTime = 0 Time = 1 min Time = t Figure 1: Tensile creep.EXPERIMENTAL PROCEDUREBecause of the high relevance to the number of creep tests that have to be performed, it is necessary to evaluate whether the samples behave in an isotropic or anisotropic manner. Standard dumb-bell-shaped specimens (ISO 527-1) have been produced in six directions, as indicated in Figure 2. These specimens were produced on a 3D Systems' Vanguard HS with layer thickness of 0.1mm, using virgin Duraform powder and standard machine build parameters.XZZXZZYYZ) a P M(s s e r t S5040302010ZYZXYXYZ XY XZInjection mouldedXXYYXY001020 Strain (% )30Figure 2: Different orientations of fabrication and their tensile behaviour (averages of five tests per direction).

The X-axis denotes the laser scanning direction, as the Z-axis denotes the layer build direction.The tensile test results reveal only little variation between XY, XZ, YX and YZ built specimens and between ZX and ZY built specimens. Therefore it was chosen to conduct creep tests only with specimens built in ZX and YX directions, representing the worst-case and best- case directions for 'short-term' properties.As a reference, injection molded Duraform specimens have been produced using an Arburg 320S 500-150 at an injection temperature of 270°C, 70 bar ram pressure leading to 525 bar effective injection pressure, 400 bar hold pressure reducing to 100 bar in 9.30 seconds. Tensile tests show that the injection molded specimens are less stiff (about 1400 MPa) and more ductile (having an elongation at break of almost 300%).6140The tensile creep behaviour has been tested according to ISO 899-1, with identical specimens as used in the tensile tests.

These specimens were connected to a frame at one side. Stresses have then been applied by attaching loads to the other side. Tests at room temperature (23°C) have been conducted with both SLS and injection molded material at approximately 12.5%, 25%, 37.5%, 50% and 80% of the material's Ultimate Tensile Strength, i.e. at 5, 10, 15, 20 and 35 MPa respectively. These tests have been conducted twice for each stress level. The extension of the specimens was measured over time by using a cathetometer. Each sequence was limited to 1000 hours.Due to the moisture uptake of the material it is of no use to conduct creep tests with dry- as-molded specimens (ISO-1874-2). It was therefore chosen to carry out the creep tests at room temperature with moistened specimens.

It was therefore chosen to carry out the creep tests at room temperature with moistened specimens. This moist state was achieved by immersing the specimens in water at 60°C for 24 hours, after which they were stored in the test environment (23°C, 50% RH) for at least 24 hours.The temperature dependent mechanical behaviour of nylon parts makes it necessary to conduct tests at a wide range of temperatures. In this study, also creep tests at elevated temperatures were conducted at high temperature (100°C) and at medium temperature (60°C). Moistening the specimens would be unnecessary, since this moist will evaporate at elevated temperatures.The elevated temperatures were obtained by testing the specimens within a heated chamber. The temperature in this heated chamber was controlled using an adjustable convection heater. Thermocouples attached to every clamp recorded the real temperature at every specimen.

The temperature in this heated chamber was controlled using an adjustable convection heater. Thermocouples attached to every clamp recorded the real temperature at every specimen. Before each test, the specimens were conditioned in the preheated environment for over 24 hours.At elevated temperatures the stress levels were set to 2.5, 5, and 10 MPa.RESULTSThe creep results are shown in Figure 3. From these curves it can be concluded that the differences between YX and ZX built specimens are negligible for small loads. These differences increase with increasing load and temperature.Raising the temperature reveals an increased initial creep response and creep rate. After 100 hours testing the 10 MPa tests show a strain of 1.7% at 23°C, 2.5% at 60°C and 3.9% at 100°C. Due to mechanical slip of clamps, the measurement of the YX specimen at 100°C had to be stopped after 1 hour.The injection molded material shows a higher initial strain and also a larger creep rate.

Using data from the material information system CAMPUS (Computer Aided Material Preselection by Uniform Standards) the creep behaviour of some general-purpose engineering plastics was reviewed. The creep behaviour of these injection molded materials, as well as the creep behaviour of the SLS processed and injection molded Duraform (IMDF) specimens is shown in Figure 4.

From these curves it can be concluded that although the SLS processed material is more sensitive to creep than most other thermoplastics it is shown to be affected less than injection molded polyamide-12.626SLS, 23°C6Injection Molded, 23°C20 MPa35 MPa55)420 MPa)4%( n i a r t S3215 MPa10 MPa%( n i a r t S3210 MPa15 MPa15 MPa0 0.010.111010010000 0.010.11101001000Log time (hours)Log time (hours)6SLS, 60°C6SLS, 100°C55)%4)%410 MPa( n i a r t S310 MPa( n i a r t S32215 MPa2.5 MPa15 MPa2.5 MPa0 0.010.111010010000 0.010.11101001000Log time (hours)Log time (hours)Figure 3: Creep behaviour of YX (), ZX () and injection molded () specimens.40Several engineering plastics, 1000h, 23°CSANPA-6PA-6 GF30PC) a P M30(s s e r t S2010PSABSSLS ZXSLS YX PA-66 IMDFPA-12000.511.522.53Strain (% )Figure 4: Isochronous curves of several engineering plastics after 1000h, 23°C.63ANALYSIS OF THE MATERIALUsing a microtome, 5µm thick slices have been planed off from the reported materials.

These slices have been applied to a cover glass and observed using a light microscope (with oblique as well as with polarised light), as can be seen in Figure 5.Figure 5: Microstructural observations. Virgin powder particle (left), average region in a ZX sample (middle left), region near the edge of the same ZX sample (middle right), injection molded sample (right).These pictures show the presence of both spherulites and kernels of particles within the SLS processed material. Near the edge of an SLS part spherulites can be found with the size of 50µm, which is equal to the size of an individual powder particle before sintering. The injection molded material shows to have smaller spherulites and no kernels of particles can be seen.Figure 6 represents the observation of a cross-section of an SLS processed sample built in YX direction. In these pictures the presence of spherulites and particle kernels can also be observed.

In these pictures the presence of spherulites and particle kernels can also be observed. These kernels show to be well-ordered according to the layered manufacturing.Figure 6: Microstructural observation of a sample built in YX direction, using a microscope with oblique (left) and polarized (right) light.64DISCUSSIONAlthough tensile tests show large differences between mechanical behaviour of YX and ZX built specimens, these differences are not reflected in the creep results. Only with creep tests at higher loads and temperatures the YX built specimens show to perform not as good as the ZX built specimens.There are two topics concerning the anisotropic behaviour of SLS Duraform material.

Influence is exerted by the weak bonding between layers [9] as well as by the presence and distribution of particle kernels within the material.The well-ordered structure of particle kernels is thought to be explained by looking at the decrease in temperature with depth in the powderbed [1]. As the laser heats up the particles the upper particles reach a temperature which causes them to melt. The temperature at the foregoing particles will be lower, so that not all particles will entirely melt.As a reference, tests have been performed with SLS processed specimens built with refreshed powder. In this material the partly-melted regions could also be found, however, the ordering of these kernels could not be demonstrated. The most liable explanation for this effect can be signed to the degradation of powder used in the reference SLS material.

The most liable explanation for this effect can be signed to the degradation of powder used in the reference SLS material. Every time the powder particles are heated they start (or continue) to degrade, which leads to an increase in melt temperature [3,5].The differences in microstructure are clarified in Figure 7, which gives a simplification of three stages, which are believed to occur in the SLS process of eight particles within one layer. The time the heat is applied and effects of shrinkage are not taken into account.Studied SLS materialReference SLS materialFigure 7: Schematics of the Selective Laser Sintering process.In the first stage, the laser passes the powder particles (represented by the dark spheres), thereby heating them to a temperature just above their melting point. During the second stage, the particles begin to melt. Gasses become enclosed within the melted material, forming pores in the material (indicated by the small spheres) [2].

During the second stage, the particles begin to melt. Gasses become enclosed within the melted material, forming pores in the material (indicated by the small spheres) [2]. Stage three is the final stage of the process. Due to the decrease in temperature with depth the upper particles show to be melted completely as the lower particles do not.65The temperature (and the time this temperature acts upon the material) in the powder bed is too low to arrange a good bonding between the processed layer and its foregoing layer. During this bonding diffusion between the layers has to take place as chains in the melt move with a snake-like motion, called reptation [9]. This random walk configuration causes entanglement between both layers. However, due to the low temperature and the short time in which the chains are able to move, the entanglement between two layers is low in respect to the entanglement of chains within a layer.

However, due to the low temperature and the short time in which the chains are able to move, the entanglement between two layers is low in respect to the entanglement of chains within a layer. Increasing the temperature can improve the entanglement between layers as is true for the injection molded Duraform. On the other hand new problems arise when increasing the temperature in the SLS process [6].Schrauwen[7] concluded that the applied temperatures and cooling rate have their effect on physical ageing and degree of crystallinity. These phenomena are most likely the main causes of the differences in creep behaviour between injection molded Duraform and SLS processed Duraform specimens.The particle kernels are thought to hinder the stretching of the material [4].

The orientation of these kernels also affects the stretching behaviour, demonstrated by the different creep behaviour of the two materials in Figure 7.The creep behaviour of the SLS material shows to be almost linear (on a logarithmic time- scale), also with large loads. Using linear extrapolation [8] creep behaviour of SLS material can be described and predicted quite exactly to a maximum of 105 hours, i.e. more than 11 years, on the hand of a 1000h test.CONCLUSIONCreep behaviour of sintered polyamide-12 processed by the SLS process has been investigated at temperatures of 23°C, 60°C and 100°C. Although the SLS processed material is more sensitive to creep than most other thermoplastics it is shown to be affected less than injection molded polyamide-12, which has been tested as a reference.Therefore, the SLS processed polyamide-12 may have the potential to be utilised for Rapid Manufacturing. Doing so, one has to keep in mind the anisotropic behaviour of the produced parts.

Doing so, one has to keep in mind the anisotropic behaviour of the produced parts. This anisotropic behaviour can be explained by looking at the relatively weak bonding between layers and the presence of partly melted powder particles. The mechanical behaviour of processed Duraform material can be explained by looking at the applied process heat, thermal history of the powder (degradation) and the cooling rate.ACKNOWLEDGEMENTThe authors acknowledge the Rapid Manufacturing Group of the LoughboroughUniversity as well as Amitek for supplying the test samples.REFERENCES[1] Bugeda, G., Cervera, M., Lombera, G, Numerical prediction of temperature and density distributions in selective laser sintering processes. Rapid Prototyping Journal, 5:1 21-26 (1999).66[2] Childs, T.H.C., Berzins, M., Ryder, G.R., Tontowi, A.E., Selective laser sintering of an amorphous polymer - simulations and experiments.

yogenic Mechanical Alloying of Poly (ether ether ketone) - Polycarbonate Composite Powders for Selective Laser SinteringJ. Schultz, R. Kander, C. Suchicital Department of Materials Science and Engineering Virginia Tech Blacksburg, VAAbstractMechanical alloying is a solid state processing technique traditionally used in the metallurgical industry to extend solubility limits in alloy systems. Mechanical alloying can also be used to blend polymer systems at ambient or cryogenic temperatures. In this work, cryogenic mechanical alloying was employed to create composite powders of Poly (ether ether ketone) sintering applications.

In this work, cryogenic mechanical alloying was employed to create composite powders of Poly (ether ether ketone) sintering applications. The (PEEK) microstructural development of the PEEK-PC system that occurs during laser sintering and the effects of this microstructure on mechanical properties of the laser sintered parts was investigated.Polycarbonate (PC)for usein selective laserKey Words: Selective Laser Sintering (SLS), Mechanical Alloying, Poly (ether ether ketone) (PEEK), Polycarbonate (PC)IntroductionMechanical alloying (MA) was first developed at The International Nickel Company, Inc. as a means of producing oxide dispersion strengthened nickel-based superalloys. This process is now a widely used fabrication method in the metals industry.

as a means of producing oxide dispersion strengthened nickel-based superalloys. This process is now a widely used fabrication method in the metals industry. Mechanical allo~ing is employed to enhance microstructural control and extend solubility limits in metal alloys[ ; however, since the late 1980s efforts have been made to apply this technology to the processing of polymer composites and blends. Research in mechanical processing of polrmers has shown improved ultimate tensile strengths of mechanically processed polymers[2 and the ability to blend immiscible polymers with no signincant degradation[3].The device used for mechanical alloying in this study. is the vibratory·mill, which provides the highest energy collisions of any milling device available. Figure 4 shows ·a. schematic representation of the motion of a vibratory mill; the arrows indicate the axis of motion of the milling vile (B).

Figure 4 shows ·a. schematic representation of the motion of a vibratory mill; the arrows indicate the axis of motion of the milling vile (B). Motion ofthe milling vile results in colHslonsbetWeenthe milling media (A).+ - ----....../r-A--- -., I , ,00 L _____ ..JI'"Bli-II.+ - ----...Figure 4: Schematic representation ofa vibratory mill.[l]311Typically, at least two balls (milling media) are used in this type of mill. Due to a decrease in the kinetic energy of the collisions with increasing number of collisions, there is a limit to the benefit of adding more balls. The vibratory mill can also be outfitted to operate at cryogenic temperatures, often required to mill polymers whose melting and glass transition temperature are near or below room temperature. Cryogenic mechanical alloying (CMA) also results in faster particle size reduction because of the increased brittleness of the polymer at low temperatures.

Cryogenic mechanical alloying (CMA) also results in faster particle size reduction because of the increased brittleness of the polymer at low temperatures. Mixing of the polymers is done when powder particles are trapped between the balls during rapid high-energy collisions with each other and the container walls. The powders are repeatedly fractured and "welded" together by the collisions, which results in the mixing of the powders and refming of the particle microstructure[4]. Figure 5 shows a schematic representation of the welding and fracture ofparticles. The particles in the A region are being welded together into a lamellar two-phase structure. The particles in region B are being fractured, thus reducing the particle size.Figure 5: Schematic Representation of high-energy collision.[l]As the milling time increases, the number of fractures and welds per particle is increased. This results in increasedrefmement of the particle morphology.

This results in increasedrefmement of the particle morphology. Figure 6 shows the progression of particle morphology refmement with increased milling time.l----l 100jJ.0'1Increased Milling Time .......Figure 6: Particle Morphology Refmement.The characteristic plate-like shape of mechanically alloyed polymers is also illustrated in Figure 6. The individual plates are made up of multiple lamellar structures of varying orientation. The312number of differentlamellar orientations in a particle increases with time, but the spacing of the lamella decreases with time[4], as shown in Figure 6.The objective of the work presented here was to demonstrate the feasibility of using cryogenic mechanical alloying to produce micro-composite powders for use in selective laser sintering. The goal of using cryogenic mechanical alloying to produce powders for SLS application is to improve the mechanical properties of parts made by SLS.

The goal of using cryogenic mechanical alloying to produce powders for SLS application is to improve the mechanical properties of parts made by SLS. Therefore, parts produced by SLS could be used as fully functional parts. The material system chosen to be cryogenically mechanically alloyed is a polymer-polymer composite of polycarbonate and poly(ether ether ketone).Experimental EquipmentTwo pieces of equipment have been designed as a part of this research, a cryogenic vibratory ball mill and a laboratory scale selective laser sintering unit. The following two sections describe the operation ofthese two devices.The Cryogenic Vibratory Ball MillMechanical alloying of the composite blend for selective laser sintering was accomplished using a cryogenic vibratory ball mill (shown in Figure 7) designed to make polymeric composite powders.

Figure 8 shows a schematic representation of the operation of the cryogenic vibratory ball mill.Figure 8: Schematic ofthe operation of the cryogenic vibratory ball mill.313A one half horsepower motor •drives the cryogenic vibratory ball mill. Th€:rotational motion of the motor is converted into linear motion by a tlywheelanda wishbone linkage that connects the primary and secondary drive shafts. Th€: milling vialt which rides on thr€:e Teflon™ guide €:ndofthemain.driveshaft. As the motorrotatest the milling railst is connected to the opposit€: vial and its contents (powdered materialandth€: mixing media) are oscillatt}dat· approximately 10 Hz through a 4.2 inch linear range of motion. During operationt the bottom inch of the cryogenic milling chamber is filled with liquid nitrogen which is continuously fed through the port in the acrylic lid. The steady-state temperature in the cryogenic mill approaches that of liquid nitrogent -196°C.

The steady-state temperature in the cryogenic mill approaches that of liquid nitrogent -196°C. Exact measurement of the temperature inside. the milling vial is not possible because of the vigorous motion impart€:d. to it. Stainless steel was chosen as the material for the milling vial and balls because its fracture toughness at cryogenic temperatures is such that the impacting surfaces do not erode· and contaminate the samples. Each run of the cryogenic vibratory ball mill produces approximately 40mL of mechanically alloyed powder.2.2 Laboratory Scale Selective Laser Sintering UnitThe selective lasersintering unit, shown in Figure 9 is a laboratory scale machine which can be used to test the applicability of a material to SLS processing using only small amounts of the material. Design and developmentofthelab-scaJe SLS unit was necessitated by the fadthatonly 40 mi of composite powder are produced for every run ofthe cryogenic vibratory ball mill.

Design and developmentofthelab-scaJe SLS unit was necessitated by the fadthatonly 40 mi of composite powder are produced for every run ofthe cryogenic vibratory ball mill. In the lab-scale SLS unit's present configurationt part geometries are limited to plaques and other flat test specimens.Figure 9: Laboratc>1J'scalei~electivelaSersinterillg l111it.The lasertmounted v€:rtically,inthe.~LSunitis>a(j()21aserwitha nominalpower oflOW. The translationofthesystemt mountedinithe bQtt9m ofthe enclosure, has a travel of 152mm x 152mm. The p.ow4er bed was· made/by Sm99thingth€: P9wder be1;We€:n1Wo<gag€: blocks with a doctor blade. Additional.layers were.ad<ied by/placingishim stock on the.ga,ge blocks and smoothing the new layer of powder with the •doctor .blade. DTM Polycarbonate was laser sintered in the lab"scale unit to provide a benchmark for the capabilities 9fthe unit and~ide the design improvements. The benchmark tests were done on ASTM D 638 Type M-II ] tensile314specimens.

The benchmark tests were done on ASTM D 638 Type M-II ] tensile314specimens. Five tensile specimens, each consisting offiye 0.25mm thick layers, were laser sintered at the following conditions: 2.6WlaserpQwer, 66mm1s scan speed, 0.2 mm scan spacing, 115°C bed temperature (obtained using an IR lamp), 3mm spot size. The average ultimate tensile strength(UTS), strainto failure,andelasticimodulus of the specimens were 3.0 ± 0.5 %and71.8±16.8MPa.\DTM\r~portsaUTS of 23 MPa[7], a strain to failure 0.6 MPa, 5.8 of 5%[7], and an elastic modulus of 1220 MPa[7] for samples produced in a Sinterstation® under standard operating conditions. The porosity of each of the tensile specimens was also determined using ASTM C 373-88[8] and the average porosity was calculated to 59 ± 2.1 %.

The porosity of each of the tensile specimens was also determined using ASTM C 373-88[8] and the average porosity was calculated to 59 ± 2.1 %. The high porQsity and low UTS relative to DTM's published value result from the fact that the lab scale SLS unit is still in the developmental stages.Procedure PEEK-PC sample productionComposite powders of 50vol% Victrex™ PEEK and 50vol% DTM Polycarbonate were made using the. cryogenic vibratory ball mill. The samples were milled for one hour and the charge ratio (mass of powder/ mass of milling media) used was 0.45. Each one-hour run produced 25.6 grams of the PEEK-PC composite powder. After milling, the PEEK-PC composite powder was laser sintered into ASTM D 638 Type M-n[6] tensile specimens.

Each one-hour run produced 25.6 grams of the PEEK-PC composite powder. After milling, the PEEK-PC composite powder was laser sintered into ASTM D 638 Type M-n[6] tensile specimens. Each specimen was 4 layers thick and was produced using the following conditions: 6 W laser power, 66 mm/s scan speed, 0.2 mm scan spacing, 125°C bed temperature, 3mm spot size, 0.25mm layer thickness.Mechanical testingFive PEEK-PC laser sintered tensile specimens, made usil1g the aforementioned conditions, were tested at 0.2 mm/s using a Texture Technologies Texture ·Analyzer load frame. Stress strain curves as well as the· UTS, elastic modulus, and strain to failure were recorded for each specimen.

Stress strain curves as well as the· UTS, elastic modulus, and strain to failure were recorded for each specimen. The average and standard deviation of the UTS, strain to failure and elastic modulus for the five specimens was calculated.Porosity testingtesting, a 25xlO mm strip was cut from each of the tensile specimens and porosity testing was performed using the Archimedes method, based on ASTM C378-88[8] with ethanol as the submersion liquid. The porosity of each specimen (the percentage of sample volume occupied by pores),wascalculated along withthe average and standard deviation for the five samples.After mechanicalScanning electron microscopyScanning electron micrographsof.tl1e as •. received.DTM Polycarbonate, cryogenically mec~apically alloyed PEEK-PC composite powder, and laserisinteredPEEK-PC·.(samesintering conditions as. above),. \V.ereta}{ellusing an International Scientific.

above),. \V.ereta}{ellusing an International Scientific. Instruments •SX-40 with.· an accelerating voltage· of 20kV.Results and DiscussionFigure 10 .shows a PEEK-PCtensilespecimenlasersinterediwith the lab~scale SItS unit. No gross distortion or geometric inaccur~ciesare visiblefto()mthephoto. The tensile specimen does, however, exhibit a large number of relatively large pores throughout the surface of the specimen.315Figure 10: PEEK-PCterisilespecimen made with the lab-scale SLS unit.Mechanical testingFigure 11 shows the stress strain curves for the PEEK-PC tensile specimens that were laser sintered using the lab scale SLS unit. The average elastic modulus, UTS and strain to failure were calculated to be 103.2±12.1 MPa, 2.8±0.4 MPa and 2.7±0.3%.

The average elastic modulus, UTS and strain to failure were calculated to be 103.2±12.1 MPa, 2.8±0.4 MPa and 2.7±0.3%. The authors believe that these low UTS and strain to failure values are a result of the large porosity observed in the laser sintered tensile specimens that results from the difficulty in achieving a smooth uniformly dense powder bed with the PEEK-PC powders. The elastic modulus of the PEEK-PC laser sintered using the lab-scale SLS is 43% higher than the elastic modulus of the DTM PC laser sintered using the lab-scale SLS unit.3.532.5:. ::IE 1 1.52(I)10.50012 Strain, %3Figure 11: Stress-strain curves for laser sintered PEEK-PC tensile specimens.Porosity testingThe porosity of the PEEK-PC s~l.1nples was quantified through porosity testing and was determined to be 54.9±2.3%. The high degree of porosity is attributed to the aforementioned difficulties in achieving a uniform dense packing of the powder bed.

The high degree of porosity is attributed to the aforementioned difficulties in achieving a uniform dense packing of the powder bed. The porous surface of a laser sintered PEEK-PC sample is showu in Figure 12.316Scanning electron microscopyFigures. 13 and 14 are scanning electron micrographs of the as-received DTM Polycarbonate composite powders·· that resulted from 1 hour of cryogenic mechanicaland the alloying.Figure 13: SEM micrograph ofDTM Polycarbonate powder.No drastic size differences are observed between the PC and PEEK-PC particles but the latter particles appear to be much rougher and more plate-like. The authors attribute. the difficulty in obtaining a smooth powder bed to the surface roughness and plate-like shape of the powder particles since the rough plate-like particles do not roll over one another as well as particles with smooth surfaces.Figure 15 shows a scanning electron micrograph of a cross sectioned surface of a PEEK..PC laser sintered sample.

Unsintered particles are visible between the dense sintered layers shown in Figure 15. The presence of these unsinteredparticlessuggests the laser sintering .parameters employed in producing this sample were inadequate in sintering through. the interlayer. Insufficient interlayer sintering.could resuI~from using a scan speed that is too high, laser power that is too low, layer thickness thatisto large, or a.

the interlayer. Insufficient interlayer sintering.could resuI~from using a scan speed that is too high, laser power that is too low, layer thickness thatisto large, or a. combination ofthese.Figure317ConclusionThe.lowUTS and strain to failure values measured and high porosity of the >PEEK-PC samples lasersintered with the lab-scaleSLS unit were attributed to difficulty in achieving awasdet~rt11inedfromthe SEM lllicr0gl'flph of the lasersintered PEEK-PC cross-section thatthe laser ~intering iparameters used<toprodllce the current samples results in insufficient interlayer sintering.Future WorkContinued development of the lab-scale SLS unit will improve the quality the samples; areas of concentration include reduction of beam spot size and improved powder delivery. These in improved mechanical properties of laser sintered improvements are expected to result samples. Investigation into methods of smoothing the rough surfaces of the PEEK-PC will also be undertaken.

Investigation into methods of smoothing the rough surfaces of the PEEK-PC will also be undertaken. other SLS material systems formed by cryogenic mechanical alloying will be ongoing.Investigation into the formationAcknowledgementsTheauthorswollid like to thankDTM Corporation for supplying the DTM<Polycarbonate. The authors·wOllld also like to thank: Dr. AlexAning for his continued support and input into our mechanical alloying research. Additional thanks goes to Julie Martin for performing the microscopy.References[llD.R..Maurice & T. H.Courtney,.Metall.Trans.,. 21,289, 1990. [2] 1.. Pan, W. J. D. Shaw. "Effects of Proce.ssing Parameters on Material Properties of[31Mechanically Processed Polyamide." Journal ofApplied Polymer Science, Vol. 56, John Wiley & Sons,. 1995. Ishida. MechanicalAlloyingofPolytetrafluorethylene with Polyethylene." Journal of Materials Science Letters, 13, 1994.[4] Farrell,M. P., R. G.

Therefore, in this study the UV curing and the infiltration behavior of the thermoset will be investigated under process relevant conditions. The investigations show a strong temperature dependent absorption of the liquid in the powder bed, whereas, the surface tension of the liquid plays a minor role. Furthermore, the UV curing of the thermosets takes place at low interaction times and at lamp powers between 100 and 300 mW/cm².Keywords: Laser sintering, thermoset, thermoplastic, multi-material partIntroductionOne trend in the field of additive manufacturing is to enlarge the applications of additive manufactured components. The range of application is limited due to the restricted material variety in powder and laser based additive manufacturing techniques.

The range of application is limited due to the restricted material variety in powder and laser based additive manufacturing techniques. Although there are efforts to develop new materials for laser sintering (LS) [1-4], the availability of polymers that can be processed in LS is still very limited compared to conventional manufacturing methods like injection molding and only single materials can be processed. Despite polypropylene and thermoplastic elastomers are launched lately, Polyamide 11 (PA11) and Polyamide 12 (PA12) hold approximately 90 % of the market share in LS [5]. The low market volume of LS powders (900 tons/year) compared to the volume of total plastic production (260 million tons/year) in combination with the high development cost of LS decelerate the launch of new polymers for LS applications [5]. Nevertheless, there is a huge customer demand for new applications of additive technologies.

Nevertheless, there is a huge customer demand for new applications of additive technologies. The local adjustment of certain properties like mechanical characteristics, stiffness or hardness and thus the generation of multi-material parts could lead to significantly higher rates of usage of LS manufactured parts, because the achieved properties could cover more specifications. In this investigation, a new hybrid additive manufacturing technique is introduced, which could overcome the gap to create new features of additively manufactured components.micro dispenserlinfrared lampmelting of SLS powder with a CO2 laser--818thermoplastic elastomer powder, is molten like in standard LS process with a CO2 laser. Temporal the injection of the liquid can be done after of before laser exposure. The order of cross-linking and melting will lead to different boundary surfaces, which shall be investigated in further studies.

The order of cross-linking and melting will lead to different boundary surfaces, which shall be investigated in further studies. After infiltration, cross-linking and melting a new powder layer is applied and the process steps start again. Like in laser sintering the thermoplastic material is cooled down after all layer are generated. The injection and curing of resins in LS process with parallel melting of the thermoplastic material lead to the generation of multi-material parts in one process.State of the artThe LS process itself can be divided in three temperature phases: preheating, building and cooling [6]. During preheating, several number of layers, which act as isolation layers, are applied in the building chamber of the LS system. The building phase consists of the three repeating steps material coating, energy input and consolidation. The building process starts with a material coating step, followed by the energy input.

The building phase consists of the three repeating steps material coating, energy input and consolidation. The building process starts with a material coating step, followed by the energy input. The cross section of the component is exposed via CO2 laser and thus one layer of thermoplastic material is molten. Finally, the building platform is lowered by the thickness of one layer, which is between 80 and 150 μm and another process cycle is introduced. During the building process the building chamber is preheated to a temperature near the onset melting temperature but above the crystallization temperature [7]. A typical time for building one layer and thus repeating the three process steps is 40 s [6]. The temperature of the surrounding powder and the exposed components decreases as the building process progresses [8].

The temperature of the surrounding powder and the exposed components decreases as the building process progresses [8]. The injection process of the reactive liquid must be realized between the material coating and the energy input or between energy input and material consolidation processes. The boundary conditions a liquid material has to fulfill to be implemented in LS process are:(cid:131)high stability at building chamber temperature(cid:131)high curing rate for low UV interaction times(cid:131)high infiltration rate, which goes along with a low viscosity (i.e. a defined surface tension between the solid and the liquid)(cid:131)sufficient bonding between the powder is indispensable(cid:131)separation of the two components may not occur.The implementation of UV reactive resins in LS process is challenging, due to the defined infiltration and curing behavior of the used material.

Furthermore, the solidification of the two materials takes place at different times. The solidification of the resin takes place in each layer and whereas the solidification of the thermoplastic occurs after the building process. Therefore, both solidification process must be investigated in detail. Nevertheless the infiltration and the curing is of main interest in this investigation and the solidification will be analyzed in further research. The curing behavior of UV curing systems can be analyzed by UV-DSC measurements, which is for example demonstrated in [9].Besides the curing of the resin, the infiltration of the liquid in the powder bed is of main importance, because it ensures a sufficient connection between two layers. The speed of infiltration and wetting of a powder layer by the used liquid does not just depend on the viscosity, but also on the surface tension (cid:592) and other characteristics.

The speed of infiltration and wetting of a powder layer by the used liquid does not just depend on the viscosity, but also on the surface tension (cid:592) and other characteristics. [10]The absorption or rather interaction of droplets with porous media is studied both experimentally and numerically in different investigations. A short overview of relevant investigations is shown in [11]. Hua Tan [11] studied the pico liter ink impact with porous media with computational fluid dynamics (CFD), validation via high speed imaging and a set up like the one described in this paper. Nevertheless, the named study is focusing on Ink Jet process with aqueous inks and paper as substrate. The interaction of a droplet with a porous media can be divided into the stages initial spreading, absorption or infiltration and evaporation.

The interaction of a droplet with a porous media can be divided into the stages initial spreading, absorption or infiltration and evaporation. The spreading kinetics is typically characterized by an equation of the following form [12]: (cid:1870)((cid:1872)) = (cid:1843) (cid:942) (cid:1872)(cid:3041)with r the time-dependent droplet radius and the constants Q and n.819(1)The infiltration or absorption can be calculated according to Denesuk et al.[13] by using the Washburn equation, assuming radial symmetry of the liquid and the porosity parameters and neglecting the finite size of the pore.

studied the spreading and infiltration behavior of small droplets (approximately 60 μm) of aqueous polymer solutions on high green density porous beds [12]. The spreading and infiltration occurs in these investigations at the same time scale, precluding their treatment as a separate phenomenon [12].The aim of this investigation is to analyze the curing behavior of UV, the curing acrylate systems and the interaction of these liquids with porous media like a powder bed. Furthermore, the boundary condition of curing and infiltration will be adapted to the temperatures in laser sintering process.Materials and MethodologyMaterialsTwo different UV curing resins (Ultra Violet Light Curing DLP Hard and Ultra Violet Light Curing DLP Flexible) from the supplier PhotoCentric 3D were used as liquids. Within these investigations, the resins were abbreviated as UV-Hard and UV-Flex. The resins were chosen due to their low viscosity.

Within these investigations, the resins were abbreviated as UV-Hard and UV-Flex. The resins were chosen due to their low viscosity. At 25 °C the viscosity of the material UV DLP Hard is given as 230 cPs [16] and the one of the material UV DLP Flexible is 200 cPs [17]. Both systems are based on an acrylate material. On the raw material, the curing behavior was analyzed by UV-DSC measurements.In laser sintering almost 90 % of the manufactured materials are based on polyamide [14]. Polyamide 12 is the plastic with the highest market share [15]. For the following investigations an unmodified virgin polyamide 12 (PA12) powder type PA 2200 from the supplier EOS GmbH, Germany is used as porous media.The curing behavior of the pure resin and the infiltration of the resin into the powder bed of polyamide 12 were analyzed to determine whether these systems can be used in combined LS process.

Therefore, the design of experiments shown in Table was conducted.TABLE I.DESIGN OF EXPERIMENTSCuring behaviorMaterialTime (s)UV power (mW/cm²)UV DLP HardUV DLP Flexible0.6, 1.2, 1.8, 2.4, 3.0, 3.6, 120 0.6, 1.2, 1.8, 120100, 200, 300100, 200, 300Infiltration behaviorMaterialTemperature (°C)UV DLP Hard / PA12 powderUV DLP Flexible / PA12 powder20 60 20 60Curing BehaviorThe curing behavior was analyzed via UV-DSC measurements. A DSC type Q2000 from TA Instruments with an UV curing unit type OmniCure S2000 from Excelitas Technologies was used. An amount of820(2)(3)approximately 5 to 10 mg of material was heated to a temperature of 50 °C. After an isothermal phase, the sample was radiated with an UV lamp and the temperature development as well as the heat flow was detected. The interaction time varies between 0.6 and 120 seconds. A second one, with an interaction time of 2 minutes, follows the first radiation.

The interaction time varies between 0.6 and 120 seconds. A second one, with an interaction time of 2 minutes, follows the first radiation. After the second radiation, the specimen should be fully cured. A schematic diagram of the UV-DSC measurement is shown in Figure 2. The temperature and the enthalpy of the first and second UV curing step was analyzed, finally.1. radiationii Q)....is':.a 0 e ~ 0., c,:s a Q) 2 ..cQ)I - - -te = 0.6, 1.2, 1.8, ... sI\2. radiation t = 2 min\time_____.FIGURE 2: Schematic diagram of an UV-DSC measurementInfiltration BehaviorFor analyzing the infiltration behavior, a specific test set up for surface tension measurements from DataPhysics was used. The measurement scheme is visualized in Figure 3. This system is usually used to determine the surface tension of liquids solids.

The measurement scheme is visualized in Figure 3. This system is usually used to determine the surface tension of liquids solids. The measuring mode sessile drop was adapted to determine the development of the drop height dependent on time.With a microliter syringe, a resin drop was applied on the powder bed surface. The drop is illuminated from one side by a light source and is observed by a CCD camera system on the other side. The infiltration of at least 10 drops was measured in order to calculate an arithmetic average. Furthermore, the minimum and maximum measurement values were determined. The development of the drop height and the contact angle was measured at 20 and 60 °C under nitrogen atmosphere.FIGURE 3: Scheme of the set up for analyzing the infiltration behavior of the resins into a powder bedResults and DiscussionCuring BehaviorTo analyze the curing behavior of the used resins UV DSC experiments were conducted.

The UV radiation power and time was changed in order to analyze the needed energy input in hybrid LS process. A diagram of these UV DSC experiments is shown in Figure 4. The figure exemplifies a measurement curve of the UV DLP hard material for laser power of 100 mW/cm² and interaction time of 0.6 seconds. For the initial radiation, the resulting enthalpy and the maximum temperature were evaluated. The second radiation is used as post curing step and therefore the radiation time is long enough to ensure that the specimen is fully cured.82180UV-Hard707065l~ 0 <;:1 1,j IJ.,) .c.6050403020101. radiation, 0.6 s2. radiation, 120 s60 55 ~ 50 ~ 63IJ.,)4540I35 -10 _ ___ ...., __ ......,,...... ____ ...., __ ......,,......_,....

radiation, 0.6 s2. radiation, 120 s60 55 ~ 50 ~ 63IJ.,)4540I35 -10 _ ___ ...., __ ......,,...... ____ ...., __ ......,,......_,.... _ _ 3000 60 120 180 240 300 360 420 480 540 600 660 720 780time (s)FIGURE 4: UV DSC measurements for the UV-Hard material for an interaction time of 0.6 seconds and power of 100 mW/cm²Figure 5 shows the UV curing DSC two factorial experiment for the system UV-Flex. The intensity of the UV lamp was changed between 100, 200 and 300 mW/cm² and the radiation time between 0.6 and 120 seconds. The maximum temperature and the enthalpy, which is released during exothermic curing reaction, is evaluated for the first and second exposure. The time for the second radiation is 120 seconds to ensure a complete curing of the samples. Therefore, with the initial radiation the reaction itself is analyzed and with the post curing (i.e. the second radiation) the residual enthalpy is evaluated.

Therefore, with the initial radiation the reaction itself is analyzed and with the post curing (i.e. the second radiation) the residual enthalpy is evaluated. With increasing time, the released energy during UV exposure rises up to approximately 300 J/g. For high radiation times the power of the UV curing step has no influence on the resulting enthalpy. However, rising the laser power from 0.6 to 1.2 seconds UV radiation time, the released enthalpy increases. The residual enthalpy, which is released during the post UV curing step decreases with increasing interaction time of the first radiation. Furthermore, the residual enthalpy is lower for higher UV powers, which is obvious due to equal total energy release of the fully cured system.T = 50 °C UV-Flex300250'.§ o 200 :>-. ~ £ 63 <.) 100 '.§ ~ 50150IJ.,)t1. exposure'Y 100 mW ♦ 200mW ◄ 300mW2. exposure'v 100 mW ◊ 200mW <l 300mW00.00.61.21.82.43.03.6120exposure time ( s)FIGURE 5: Enthalpy vs.

exposure'Y 100 mW ♦ 200mW ◄ 300mW2. exposure'v 100 mW ◊ 200mW <l 300mW00.00.61.21.82.43.03.6120exposure time ( s)FIGURE 5: Enthalpy vs. UV power and radiation time for the first and second radiation of the UV-Flex materialThe maximum temperatures during UV exposure for the material UV-Flex is represented in Figure 6. During first exposure, the maximum temperature of the material rises with increasing interaction time, which goes along with the curing progress and the released energy. Furthermore, the energy input has no significant influence on the resulting maximum temperature during UV irradiation. The temperature of the sample during the second heating is lower than for the first UV exposure with a value between 50 and 60 °C.822100T = 50 °C UV-Flex90G~ ll)~ 63 I8070605040~:! • ~ .. I1. exposure~.... l00mW ♦ 200mW ◄ 300mWIi2.

I1. exposure~.... l00mW ♦ 200mW ◄ 300mWIi2. exposurel00mW 'v ◊ 200mW <I 300mWT !<I300.00.61.21.82.43.03.6120exposure time ( s)FIGURE 6: Maximum temperature in UV DSC vs. UV power and radiation time for the first and second radiation of the UV-Flex materialThe residual enthalpy shown in Figure 5 is for the second radiation lower than for the first one. Thus, the temperature development due exothermic reaction is reduced. Furthermore, the maximum temperature is independent from the energy input and it is decreasing slightly with increasing exposure time. For long impact time and thus a high degree of cure or polymerization, nearly no increase of temperature can be detected. Compared to the basal temperature of 50 °C the specimen’s temperature is not heightened for long radiations times during the second heating.

Compared to the basal temperature of 50 °C the specimen’s temperature is not heightened for long radiations times during the second heating. For further investigations the basic temperature will be increased in order to analyze whether there is an influence of the surrounding temperature on the cross linking process. The aim would be to simulate the process conditions and therefore temperature up to 150 °C (assumed interaction temperature of the resin with a PA12 powder, which was manufactured at a building chamber temperature of 167 °C) will be investigated.Figure 7 and Figure 8 shows the enthalpy and the maximum temperature of the UV DSC measurements for the first and second irradiation for the UV-Hard material. The enthalpy is rising with increasing UV radiation time and with increasing UV power during the first heating. This trend could also be observed for the material UV-Flex. The absolute level of released energy is higher for the material UV-Flex compared to UV-Hard.

This trend could also be observed for the material UV-Flex. The absolute level of released energy is higher for the material UV-Flex compared to UV-Hard. The evaluation of the second radiation shows a reduction of the enthalpy with increasing time of the first UV exposure. For high impact times the residual enthalpy tends to zero. Thus, the whole reaction takes place during the first exposure. To adapt this behavior to the hybrid laser sintering process it is important to ensure that a residual amount of material can react during the second UV irradiation to get a sufficient connection between two applied layers of UV resin. Therefore, short exposure time are favored for the final hybrid process.T = 50 °C UV-Hard300250~ ;:::; ';: 200 ~1 150 ~ ~ 50(.)100ll)0I i i II I!1:i_~ = •I:Q: :a:Q: A== •Iill- 6s- D.:0: I1. exposure■ l00mW200mW.& 300mW2.

exposure■ l00mW200mW.& 300mW2. exposurel00mW D 0 200mW 6 300mW0.00.61.21.82.43.03.6120exposure time ( s)823The resulting maximum temperature during the first and second UV irradiation is represented in Figure 8. With rising impact time, the temperature during first irradiation increases from approximately 55 °C (100 mW/cm²) to 65 °C (100 mW/cm²). Furthermore, the temperature seems to be dependent on energy input. For higher UV power the maximum temperature rises. The post curing after the initial UV curing steps shows an increase of temperature for higher radiation times. Due to a lower value for exothermic reaction, the temperature rise is reduced and for the specimen with a radiation time of 1.8 seconds (first radiation) no significant temperature increase compared to the basal tempering could be detected. In comparison to the material UV-Flex the temperature value of the resins UV-Hard is lower. T = 50 °C UV-Hard 1.

In comparison to the material UV-Flex the temperature value of the resins UV-Hard is lower. T = 50 °C UV-Hard 1. exposure10090G 80~ ll)70~ 63 I 60 ~ 5040■ l00mW • 200mW ... 300mW T I I I i};:]I:■:: 2. exposure.J.TJ:._l00mW D 0 200mW /':, 300mW• II I♦ ~ ~ -ill-'°''°'.,=±=i300.00.61.21.82.43.03.6120exposure time ( s)FIGURE 8: Maximum temperature in UV DSC vs. UV power and radiation time for the first and second radiation of the UV-Hard materialInfiltration BehaviorFor combined LS process with reactive resins the curing and the infiltration behavior plays a major role. Therefore, the infiltration was analyzed with a test set up, which was adapted from surface tension measurements. Figure 9 shows the development of droplet height for two different surrounding temperatures for the material UV-Flex. As substrate a PA12 powder was used.

Figure 9 shows the development of droplet height for two different surrounding temperatures for the material UV-Flex. As substrate a PA12 powder was used. The median curve was calculated on basis of at least 10 droplets and the black and red lines show the minimum and maximum values of the measurements. With increasing temperature, the infiltration of the liquid in the powder bed is accelerated.2.0,.-..,!1.5 "Eh ' [) ~ 1.0 0. ~+-'UV-FlexT = 20 °C 0 T = 60°C - - - - · min. values - - - - · max. values0.50.005152025time (s)FIGURE 9: Droplet height vs. interaction time and temperature for the UV-Flex materialFor a temperature of 20 °C the infiltration takes approximately 23 seconds, whereas for 60 °C the absorption is finished after approximately 2 seconds. Measurements for higher temperatures near building chamber temperature cannot be realized with this test set up, due to short infiltration times.

Measurements for higher temperatures near building chamber temperature cannot be realized with this test set up, due to short infiltration times. Therefore, the set up824will be equipped with a high speed camera in the future. For average layer time in LS of approximately 40 seconds the infiltration process is fast enough to avoid a slowdown of the whole production process. The initial droplet height is independent from the surrounding temperature.The droplet development for the material UV-Hard is shown in Figure 10. As substrate material, a polyamide 12 powder with a bulk density of 0.43 g/cm³ was used. The infiltration for such absorption process shows an equal trend than the one for the UV-Flex material (Figure 9). The infiltration depends on the temperature, which can be traced back to the reduction of viscosity with increasing temperature.

The infiltration depends on the temperature, which can be traced back to the reduction of viscosity with increasing temperature. The initial droplet height shows a similar level for both analyzed materials.Link in curing experiments in further investigations the temperature of the powder bed will be increased to simulate the laser sintering process. In laser sintering of PA12 the building chamber is heated up to 167 °C and thus assumed interaction temperatures lay between 140 and 160 °C when the droplets have a temperature of 50 or 60 °C. Unfortunately, with the shown set up the infiltration for higher temperatures cannot be traced. Therefore the set up is equipped with a high speed camera in the future.2.0,-.,],1.5 ..... ...c:: 00 ·v :: 1.0UV-HardT =20 °C □ T = 60 °C - - - - · min. values - - - - · max. valuesQ)l 0.50.00510152025time (s)FIGURE 10: Droplet height vs.

...c:: 00 ·v :: 1.0UV-HardT =20 °C □ T = 60 °C - - - - · min. values - - - - · max. valuesQ)l 0.50.00510152025time (s)FIGURE 10: Droplet height vs. interaction time and temperature for the UV-Hard materialSummary and OutlookIn this research work, a new process route to produce multi-materials parts in laser sintering (LS) is introduced. The process itself is adapted from powder binder techniques combined with laser sintering of polymers. Therefore, reactive resins were implemented in these LS processes. Besides the melting of the thermoplastic materials, a resin is cured via UV light. The curing behavior of UV acrylate based resins and the infiltrations performance of the liquid in combination with a powder bed were analyzed.The curing of the materials was analyzed via UV DSC measurements and under variation of impact time and UV power. With increasing time and power the exothermic reaction progresses.

With increasing time and power the exothermic reaction progresses. Furthermore, the temperature of the specimen rises due to higher exothermic reaction. The specimens were exposed to the UV light source a second time in order to analyze the residual enthalpy. For long initial curing times no significant energy release could be detected, which indicates a complete curing during the first UV light exposure. This time and power combination are not favored due to the need of bonding between two layers, which goes along with a residual reaction energy.Besides the curing, the infiltration is analyzed with a specific test set up, which is adapted from surface tension measurement. The infiltration time for both resins into a PA12 powder bed take less than 25 seconds at room temperature. For an expected layer time of 40 seconds, an infiltration speed lower than 25 seconds is favored. With increasing temperature, the infiltration of the dispended droplet is even faster.

For an expected layer time of 40 seconds, an infiltration speed lower than 25 seconds is favored. With increasing temperature, the infiltration of the dispended droplet is even faster. Thus, both materials will infiltrate the powder bed in an adequate time. Nevertheless, the parallel running of the curing and infiltration steps should be analyzed in further studies. Therefore, the test set up will be modified with an UV light source.825Furthermore, the curing and the infiltration will be analyzed at higher temperatures, which are nearer at the laser sintering of thermoplastic materials in further investigations. For PA12 powder the building chamber temperature is for example between 165 and 172 °C. The droplets will be stored in a tempered syringe at a temperature between 50 and 60 °C. Thus, a temperature between 140 and 160 °C is assumed as interaction temperature.

The droplets will be stored in a tempered syringe at a temperature between 50 and 60 °C. Thus, a temperature between 140 and 160 °C is assumed as interaction temperature. Therefore, both test set ups will bed adapted in the future to analyze the curing and infiltration behavior process adapted.AcknowledgementFunded by the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation) – Project-ID61375930 – SFB 814, sub-project B7.References[1] Schmidt J, Sachs M, Zhao M, Fanselow S, Wudy K, Drexler M, et al. A novel process for production of spherical PBT powders and their processing behavior duringlaser beam melting. AIP Conference Proceedings. 2016;1713:140008.[2] Wegner A, Harder R, Witt G, Drummer D. Determination of Optimal Processing Conditions for the Production of Polyamide 11 Parts using the Laser SinteringProcess. International Journal of Recent Contributions from Engineering, Science & IT (iJES). 2015;3:pp.-5-12.[3] Amado A, Wegener K, Schmid M, Levy G.

ring Behavior of Ceramic Resin for StereolithographyG. Allen Brady, Tien-Min Chu and John W. Halloran Materials Science and Engineering Department, College of Engineering The University of Michigan, Ann Arbor, MI. 48109ABSTRACTCeramic green bodies have been created with stereolithography methods by using ultraviolet cur able suspension of ceramic powders, a "ceramic resin", on a Stereolithography Apparatus (SLA). A minivat system with mini-recoater blade was designed to run small batch experiments on an SLA-250/40 machine. With the ceramic resins developed from hydroxyapatite for biomedical ap plication and from alumina and silica for making metal casting molds, ceramic green bodies were built. Diagnostic parts of these ceramic resins were made with the Accumax diagnostic kit (3D Systems) and the curing parameters determined.

Diagnostic parts of these ceramic resins were made with the Accumax diagnostic kit (3D Systems) and the curing parameters determined. The effect of shrinkage stress and scattering will be discussed.IntroductionSince the early 1990's, the Stereolithography (SL) technique has been the predominant fab rication method for rapid prototyping[l]. The SL technique involves solidification of a liquid resin by a UV-laser source. The laser cures one layer by "drawing" it on the surface of a vat of resin. The 3-dimensional part is built layer by layer. The research at the University of Michigan is fo cused on developing new materials, particularly ceramic materials, for use in a Stereolithography Apparatus (SLA)a. These ceramic resins are used in the same manner as the traditional epoxy or acrylate resins. However, instead of producing a 100% polymer part, these ceramic resins yield a ceramic green body consisting of approximately 50% ceramic and 50% uv-polymerized binder.

However, instead of producing a 100% polymer part, these ceramic resins yield a ceramic green body consisting of approximately 50% ceramic and 50% uv-polymerized binder. Current materials include alumina, silica and hydroxyapatite (HA). Potential applications for the alumina and silica resins include cores and shell molds for investment casting as well as functional ceramic prototypes. HA ceramics have significant biomedical applications[2,3].These ceramic resins are fluid suspensions of ceramic particles in a uv-curable medium. For ceramic processing purposes, the resins must be highly loaded (at least 50 volume%) to pro duce reliable, dense finished ceramic parts.

For ceramic processing purposes, the resins must be highly loaded (at least 50 volume%) to pro duce reliable, dense finished ceramic parts. For SL processing purposes, the resins must have a low viscosity « 5000 cps) to undergo the recoating processb, and self-leveling.The cure depth is defined in the "Working Curve" equation,(1)where E is the average exposure dose supplied by the laser (controlled by scanning speed, vs)' Cd, is the user definable cure depth and Dp and Ec are resin parameters respectively known as deptha. Model SLA250/40, 3D Systems, Valencia, CA. b. Recoating here refers to the doctor-blade type of recoating found on SLA 250403of penetration and critical exposure. A schematic of resin curing under a laser is shown in Figure 1. While intensity attenuation for traditional resins used for SL is limited by absorption, the inten sity decrease of the uv-radiation in these turbid suspensions is controlled by scattering.

While intensity attenuation for traditional resins used for SL is limited by absorption, the inten sity decrease of the uv-radiation in these turbid suspensions is controlled by scattering. The fol lowing relation has been used to describe the effect of scattering on cure depth[3,5],(2)where <l> is the volume fraction solids, d is the particle size and t1.n is the difference in index of re fraction between the suspended particles (ceramic) and medium (photocurable monomer resin). The ~ term involves interparticle spacing and wavelength of radiation.Materials and MethodsThree resins systems are examined in this work, alumina-aqueous, silica-acrylate and hy droxyapatite-acrylate. Details of resin fabrication are given elsewhere[3]. The alumina-aqueous resin is 52vol% solids with a monomer solution consistingC of 27% N-hydroxymethacrylamide, 3% methylene bis-acrylamide, 35% water and 35% ethylene glycol.

The alumina-aqueous resin is 52vol% solids with a monomer solution consistingC of 27% N-hydroxymethacrylamide, 3% methylene bis-acrylamide, 35% water and 35% ethylene glycol. The silica-acrylate resin has 43vol% solids, and the monomer solution is 30% 2-ethoxyethoxy ethyl acrylate and 70% hexane diol diacrylate. The HA-acrylate resin has 40vol% solids and the monomer solution is 66% octyl decyl acrylate, 22% hexane diol diacrylate and 12% trimethylolpropane triacrylate.All parts were produced on an SLA 250/40 with a He/Cd laser source. Typically, the laser power at the surface of the resin was about 30 mW. The machine was equipped with an exchange able minivat (lL to 2L) with a recoating arm (shown in Figure 2) for small sample experiments. Some single layer parts were produced in a glass petri dish on an adjustable stand rather than the minivat.

Some single layer parts were produced in a glass petri dish on an adjustable stand rather than the minivat. The Windowpane technique[l] was used to measure the cure depth as a function of ex posure from which the resin parameters, Dp and Ec were determined for each resin. After curing in the SLA-250, parts were postcured 30 minutes in a uv chamber and allowed to dry at ambient temperature.. To remove the photopolymers from the postcured parts, a burnout was performed in air. The parts were heated slowly (l°C/min) to 500°C, held 1 hour, then cooled. To sinter the alu mina parts, they were heated in air to 1550°C (10°C/min), held 1 hour then cooled.ResultsThe cure depth as a function of energy dose is plotted in Figure 3 for the ceramic resins and the standard epoxy resin. In each case, the cure depth is linear with In(E), although the slopes(Dp) do differ widely. The resin parameters were determined from these data and are listed in Table 1.

In each case, the cure depth is linear with In(E), although the slopes(Dp) do differ widely. The resin parameters were determined from these data and are listed in Table 1. In each case, the linear relation between the cure depth and In(energy dose) is apparent. The effect of these resin parameters are shown qualitatively in the cured line profiles shown in Figure 4 for silica and alumina. As the Dp decreases, the cured profile becomes wider and more shallow. A "Block M" shape was made in alumina and HA and also in SL 5170 for comparison. An epoxy M, an alumina M after post-curing and an alumina M after sintering are shown in Figure 5 (from right to left respectively). From this photo, curing and drying shrinkage of the alumina-aqueous resin is not much larger than that of the SL 5170. The significant portion of shrinkage occurred in the sin-c. all percentages are given in weight percent unless otherwise noted.404tering step, approximately 12% (linear).

The significant portion of shrinkage occurred in the sin-c. all percentages are given in weight percent unless otherwise noted.404tering step, approximately 12% (linear). No cracking or significant warping occurred during either curing, drying, or sintering. A fracture surface of one section of the alumina M is shown in Figure 6. The image to the left shows a scalloped edge resulting from the wide, shallow shape of the cured line profile. At higher magnification in the right image, the interface between two layers appears to have sintered very well with no remnants of the layered processing evident.Table 1. Resin ParametersResinDp, JlmEc' mJ/cm2Cibatool® SL-517011910.2Al20 3-Aqueous6919.4SiOTAcrylate26013.4HA-Acrylate5610.8A portion of a the Block M shape built from HA-acrylate is shown in Figure 7. The side surface also has a scalloped appearance, but the layers appeared to be knit together well.

The side surface also has a scalloped appearance, but the layers appeared to be knit together well. However, shrinkage cracking occurs during curing and is shown in Figure 8 to occur between the borders on the layers.DiscussionAll resins used in this work follow a predictable relation between cure depth and energy dose. Work is continuing to increase the stiffness and solids loading of these ceramic resins. Layer cure depths of 200 Jlm to 600 Jlm are achievable with the silica resin, 180 Jlm to 300 Jlm for the alumina resin and 200 Jlm to 300 Jlm for the HA resin. Parts have been built with adequate layer ing. The microstructure of a part fabricated in the alumina resin reveals no remnants of layered proceSSIng.Side scattering produces the scalloped edges found on some initial parts built at Michigan. A side view of several cured lines drawn one upon another are shown in Figure 9 which illustrates the scalloping effect. The phenomenon also occurs in alumina-aqueous resins.

A side view of several cured lines drawn one upon another are shown in Figure 9 which illustrates the scalloping effect. The phenomenon also occurs in alumina-aqueous resins. The refractive in- dex difference (.&n2 ) in the alumina-aqueous is approximately 0.076. Reducing Lln2 by using an acrylate-based solution (n= 1.447 vs. 1.425 for aqueous solution) would reduce Lln2 to 0.064, sub sequently decreasing the level side-scattering while increasing the cure depth.Shrinkage cracking occurred in the acrylate resin systems (with both silica and HA). This cracking did not occur in the alumina-acrylate system which only experienced minor shrinkage «3% linear). There are two simple ways to alleviate these shrinkage stresses. One is to employ an unreactive diluent in the acrylate solution. This decreases the concentration of the acrylate bonds in the resin and thus the curing shrinkage.

One is to employ an unreactive diluent in the acrylate solution. This decreases the concentration of the acrylate bonds in the resin and thus the curing shrinkage. For example, in the alumina-aqueous system, wa ter and ethylene glycol are unreactive diluents which dry after curing. The second method is to make use of new build styles. Build style refers to the algorithms used to generate the path of the laser when drawing each layer. The power of optimizing the build style to control shrinkage is seen405in the use of SL5170 epoxy resin with the ACESTM build style. The theoretical shrinkage is 6.2%, but use of this build style manages the shrinkage to an effective 0.05% in the x and y directions! Experimentation on build styles for the specific ceramic resins is currently underway.Scattering contributes to lower depths of penetration by the uv-radiation as well as to larger line widths.

The overall shapes of the line profiles are more shallow than in the unfilled resin also due to scattering. As a result, more radiation is delivered in a sideways direction which increases the line width. In order to build accurate parts, this increased line width must be dealt with in the build styIe. Work is underway to characterize and model the shape of the cured line profile for these resins. Despite the progress made to date on elementary fabrication, shrinkage issues and the effect of scattering emphasizes the need to fully exploit the capabilities of the SLA-250 by creating new build styles to both accommodate the modified line profile and alleviate shrinkage and distor tion.AcknowledgmentThe research on the refractory ceramics is supported by the Office of Naval Research under contracts N00014-95-1-0527 through Dr. S. Fishman and R. Wachter and the research on biocer amics is supported through the University of Michigan by a Rackham Grant to T-M. Chu.References 1.P.

M. L. Griffith, "Stereolithography of Ceramics", Ph.D. Thesis, University of Michigan, March, 1995. M. L. Griffith and J. W. Halloran, Stereolithography", submitted to Journal of the American Ceramic Society.2.3.4.5."Free Form Fabrication of Ceramics by406FIGURESzFigure 1 Schematic of a cured line profile generated by laser scanning on the resin surface.Figure 2 Photos of exchangeable minivat system. Elevator platform and recoating blade are shown at the right.4071000 . , - - - - - - - - - - - - - - - - - - - - - - ..

Note scalloped edges on left. Close-up picture on the right shows no evidence of layering.Figure 7 Tilted vie",- of HA-acrylate "1\'1" showing scalloped edges. Note contrast between border and fill regions.409Figure 8 Shrinkage cracks in HA-acrylate between borders.Figure 9 Stacking of cured profiles in HA-acrylate resin.410.

Unfortunately, the SLS process is restricted regarding the materials that can be processed and thus resulting component properties are limited.The investigations in this study illustrates a totally new additive manufacturing process which combines reactive liquids like thermoset resins and thermoplastics to generates multi material SLS parts. To introduce thermoset resins into the regular SLS process, the time- temperature-dependent curing behavior of the thermoset and the infiltration behavior has to be understood in order to assess the process behavior. The curing behavior was analyzed by rotational viscosimeter. Furthermore, the fundamental infiltration behavior was analyzed with micro dosing infiltration experiments. Finally, a thermoset resin in combination with a dosing system was choosen for integration in a laser sintering system.Introductionlaser sintering (SLS) of thermoplastic powders, generate components directly from a CAD dataset without needing a form or a mold.

Thus, the layerwise process of selective laser sintering allows the generation of individualized complex serial parts. Contrary to conventional manufacturing techniques like injection molding, which are optimized for high volume production and low unit costs, the costs for additive manufactured parts are almost independent from the degree of complexity and the quantity [1]. The advantages of additive technologies are favored for products with high customization, a high level of complexity and small lot sizes, which is known as “mass customization” [2].Additive manufacturingtechnologies,like selectiveAlthough there are efforts to develop new materials for SLS [3-7], the availability of polymers that can be processed in SLS is still very limited compared to conventional manufacturing methods like injection molding.

Beside polyamide 12 and polyamide 11, filled systems and a few other thermoplastic materials like polypropylene and thermoplastic elastomers are available at the market [8]. The reason for the slow development of new polymers powders for SLS is the small market value. The market for SLS powders is with 900 tons/year compared to 260 million tons of worldwide plastic used very small [8]. Nevertheless, there is a huge customer demand for new applications of additive technologies. The limited component characteristics, which can be reached with the available materials, hinders the use of SLS for specific applications. The local adjustment of certain properties and thus the generation of multi-material parts could lead to significantly higher rates of usage of SLS manufactured parts, because the achieved properties could cover more specifications.

Therefore, in this paper a new unique additive manufacturing process, which overcome the bridge to create new properties in SLS, will be introduced. The hybrid process between powder328binder techniques and SLS integrates reactive liquids in the SLS process, as shown in Figure 1. These fluid materials are planned to be injected into a conventional SLS powder bed through a micro valve system and the reaction of the liquid is initiated via IR or UV radiation. Beside the reactive liquid, the thermoplastic powder can be processed in SLS like usually. Thus, multi material parts consisting of a reactive system like an epoxy resin and a thermoplastic material can be generated in one process.Figure 1: Schematic visualization of the hybrid SLS process with reactive liquidsState of the artThe SLS process itself can be divided in three temperature phases: preheating-, building- and cooling phase [9].

The building process consist of the three steps: material coating, energy input and consolidation. Firstly, a layer of powder is applied in the building chamber. The layer thickness of the coated layer is usually between 80 and 150 µm. Secondly the cross section of the component is exposed by a CO2 laser and the materials starts to melt selectively. Finally, the building platform is lowered by the thickness of one layer and the three process steps will be repeated. During building process the building chamber is preheated to a temperature near the onset melting temperature but above the crystallization temperature [10]. A typical time for building one layer and thus repeating the three process steps is 40 s [9]. The temperature of surrounding powder and the exposed components decrease the farther away the point of measurement is from the powder surface [11].

The temperature of surrounding powder and the exposed components decrease the farther away the point of measurement is from the powder surface [11]. The implementation of the reactive liquid in the hybrid process must be done after material coating and before or parallel to the energy input via CO2 laser. The boundary conditions a liquid material has to fulfill to be implemented in SLS process are: high stability at building chamber temperature  specific curing behavior under the temperature occurring in SLS  high infiltration rate, which goes along with a low viscosity a defined surface tensionbetween the solid and the liquid separation of the two components may not occur.Thus introducing a material that reacts during SLS process presents a specific challenge. In order to develop thermosets for SLS process, the curing under a specific temperature profile, which is determined by the SLS process itself, must be monitored precisely.

In order to develop thermosets for SLS process, the curing under a specific temperature profile, which is determined by the SLS process itself, must be monitored precisely. The thermal household during SLS will be influenced for example by the energy input (laser power, scan speed, …), the number of exposed components, the chamber and feeder temperature. Besides the curing of the resin the infiltration of the liquid in the powder bed is of main importance, due to ensure a sufficient connection between two layers. The speed of infiltration and wetting of a powder layer by the used liquid does not just depend on the viscosity, but also on the surface tension σ and other characteristics, as exemplarily depicted in Figure 2. [12]329Figure 2: Fluid powder bed interactionThe absorption or rather interaction of droplets with porous media is studied both experimentally and numerically in different investigations. A short overview of relevant investigations is shown in [13].

A short overview of relevant investigations is shown in [13]. Hua Tan [13] studied the pL impact with porous media with computational fluid dynamics (CFD) and validation via high speed imaging and uses a similar set up like described in this paper. Nevertheless, the named study is focusing on Ink Jet process with aqueous inks and paper as substrate. The interaction of a droplet with a porous media can divided into the stages initial spreading, absorption or infiltration and evaporation. The spreading kinetics is typically characterized by an equation of the following form [14]:𝑟(𝑡) = 𝑄 ∙ 𝑡𝑛(1)with r the time-dependent droplet radius and the constants Q and n.The infiltration or absorption can be calculated according to Denesuk et al.[15] by using the Washburn equation and assuming radial symmetry of the liquid and the porosity parameters and neglecting the finite size of the pore.

The infiltrated volume VP at time can be calculated with the following equation:𝑉𝑝(𝑡) =𝑘2𝑡 ∫ 0𝑟2(𝑡′) √𝑡′ 𝑑𝑡′(2)𝑘 = 𝜋 ∙ 𝑎𝑝√𝛾𝐿𝑉∙𝑐𝑜𝑠𝜃∙𝑅𝑝𝑜𝑟𝑒 2𝜂(3)with Rpore the pore radius, ap the area fraction of pores, 𝛾LV the surface tension of the liquid, 𝜃 the contact angle between the liquid and solid and 𝜂 the viscosity.Holman et al. studied the spreading and infiltration behavior of small droplets (approximately 60 µm) of aqueous polymer solutions on high green density porous beds [14]. The spreading and infiltration occurs in this investigations at the same time scale, precluding their treatment as separate phenomena [14].The aim of this investigation is to analyze spreading and infiltration behavior with a specific test set up and determine whether spreading and infiltration can be separated.

Furthermore, the influence of the pore size or rather bulk density as well as the surface tension of the used liquid on the infiltrations behavior is main interest.MethodologyMaterial / powderFor the following investigations an unmodified polyether block polyamide powder (PEBA) type PrimePart ST from the supplier EOS GmbH, Germany is used. PEBA is a thermoplastic elastomer consisting of polyamide and polyether backbone blocks. Two different aging states of the powder, virgin and used powder was used, in order to analyze the infiltration330behavior dependent on the bulk density of the material. Figure 3 shows a scanning electron microscopy image of the PEBA powder.Figure 3: Scanning electron microscopy image of PEBA powderA DSC diagram of the used PEBA is shown in Figure 4. On basis of the DSC analysis, the building chamber temperature should be chosen between 125 and 137 °C for this material.

On basis of the DSC analysis, the building chamber temperature should be chosen between 125 and 137 °C for this material. Furthermore, material properties of the used resins like surface tension and viscosity are of main interest at a temperature interval near building chamber temperature. Due to reduction of temperature during coating machismo of up to 30 K. The relevant temperature range is between 100 and 130 °C.Figure 4: DSC diagram of the used PEBA powderMaterial / resinsAs reactive liquid two epoxy resins Araldite GY 764 and Araldite GY 793 purchased from Huntsman Advanced Materials, Switzerland was used. Araldite GY 764 is based on bisphenol A, whereas GY 793 is based on bisphenol A/F. The used curing agent, Hardener XB 3473, was also purchased from Huntsman Advanced Materials. The used mixing ratios were 100:23 parts per weight for GY 764 and 100:22 parts per weight for GY 793.

The used curing agent, Hardener XB 3473, was also purchased from Huntsman Advanced Materials. The used mixing ratios were 100:23 parts per weight for GY 764 and 100:22 parts per weight for GY 793. The epoxy resins were selected because of their low viscosity and their high reaction temperatures, which deem them compatible with the sought processing through a micro valve in the hot building chamber. GY 764 possesses a viscosity of 350 - 550 mPas, GY 793 650 - 750 mPas and the Hardener XB 3473 80 - 125 mPas [16]. For the further investigations, the mixtures of Araldite and Hardener are abbreviated with the pure resin names Araldite GY 764 and Araldite GY 793.331Characterization of material propertiesThe surface tension of the used resins was analyzed with the pendant drop method and the surface tension measurement set up from the company DataPhysics. At least 10 drops were dosed and evaluated according their dimensions.

At least 10 drops were dosed and evaluated according their dimensions. The Laplace-Young equation was used to calculate the surface tension of the liquid.The viscosity of the resins was determined with a rotational viscometer type Discovery HR-2 from TA Instruments. As measurement geometry, a plate-plate set up was used. The frequency was set to 0.1 Hertz and the normal force was 10 Pascal. The uncured mixture was applied between the two plates at a starting temperature of 25 °C and the specimen was heated with a heating rate of 2 K/min up to 200 °C. Finally, the complex viscosity was analyzed dependent on the temperature.Beside the influence of the temperature, the influence of the bulk density on the infiltration behavior will be investigated. Therefore, the bulk density was analyzed according to DIN EN ISO 60 [17]. The bulk density is defined as the mass of the polymer particles divided by the total volume they occupy.

Therefore, the bulk density was analyzed according to DIN EN ISO 60 [17]. The bulk density is defined as the mass of the polymer particles divided by the total volume they occupy. For the analysis a bulk density tester from Emmeran Karg Industrietechnik type ADP is used. The average was calculated out of at least five measurements.Infiltration behaviorFor combined laser sintering process the infiltration behavior of the reactive liquid in the powder bed is of main interest. Therefore, the infiltration behavior was analyzed with a specific test set up, which is schematically illustrated in Figure 5. For the measurements, a basic set up from the company DataPhysics was used. This system is usually used to determine the surface tension of liquids with the pendant drop method and the surface tension of solid with for the sessile drop method. The measuring mode sessile drop was adapted to determine the development of the drop height dependent on time.

The measuring mode sessile drop was adapted to determine the development of the drop height dependent on time. Thus, the infiltration of the liquid in the powder bed can be evaluated. Furthermore, the angle between the powder bed surface and the liquid was analyzed.With a microliter syringe, a resin drop was applied on the powder bed surface. The drop is illuminated from one side by a light source and is observed by a CCD camera system on the other side. The infiltration of at least 10 drops was measured in order to calculate an arithmetic average. The development of the drop height and the contact angle was measured at 20, 40 and 60 °C under nitrogen atmosphere. Figure 4 shows, that the temperature range between 100 and 130 °C is of main interest due to building chamber temperature. Nevertheless, the infiltration at these temperature takes less than 1 second and is hardly analyzable.

Nevertheless, the infiltration at these temperature takes less than 1 second and is hardly analyzable. Table 1 represents the detailed design of experiments.Figure 5: Scheme of the measurement of the infiltration behavior of the resin in a powder bed Table 1. Design of Experiments332ResinBulk density powder (g/cm³)Temperature (°C )Araldite GY 764Araldite GY 7930.390.360.390.3620 40 60 20 40 60 20 40 60 20 40 60Results and discussionFirst, the bulk density of the used substrate materials, a thermoplastic elastomer with a volumetric median particle size of 77 µm, was analyzed. As substrate material virgin powder and powder that has passed through one processing cycle was chosen. The virgin powder shows an average bulk density of 0.39 g/cm³, Figure 6. With increasing number of processing cycles, the bulk density decreases to a value of 0.36 g/cm³, Figure 6.

The virgin powder shows an average bulk density of 0.39 g/cm³, Figure 6. With increasing number of processing cycles, the bulk density decreases to a value of 0.36 g/cm³, Figure 6. The packing density of the bulk material can be calculated by dividing the bulk density and the density of the solid material. Thus, the packing density of the used material is 35 % and 33 % for a solid material density of 1.1 g/cm³.Figure 6: Bulk density of the substrate powderBeside the influence of the porosity of the substrate material, the influence of the used liquid resin and thus the surface tension and it’s viscosity on the infiltration behavior was determined. Therefore, the surface tension of the resins Araldite GY 764 and Araldite GY 793 was measured via pendant drop method and calculated according the Laplace-Young equation, Figure 7. The resin with the trade name Araldite GY 764 shows with a value of 37 Nm/m a higher surface tension than the resin with the trade name Araldite GY 793.

The resin with the trade name Araldite GY 764 shows with a value of 37 Nm/m a higher surface tension than the resin with the trade name Araldite GY 793. Nevertheless, the surface tension cannot be linked directly to the infiltration speed due the influence of viscosity and contact angle on the infiltrated volume expressed in equation 3.333Figure 7: Surface tension of the used resinsThe viscosity and surface tension are determining the absorption behavior of liquids in porous structures according to equation 2. For this reason, the temperature-dependent viscosity of the resins was analyzed. Figure 8 shows the complex viscosity in dependency of temperature for a heating rate of 2 K/min. With increasing temperature, the viscosity shows a first decrease, which can be linked to the higher mobility of chain segments with rising temperature. For higher temperatures, the crosslinking or rather curing process takes place and the viscosity rises.

For higher temperatures, the crosslinking or rather curing process takes place and the viscosity rises. The minimum viscosity is reached in a temperature interval between 100 and 120 °C for both resin and hardener mixtures. The curing reaction starts for a heating rate of 2 K/min at a temperature of 100 °C. This value depends is time and temperature and thus the heating rate plays a major role. Comparing the two resins, the curing behaves almost equal. The system with the trademark Araldite GY793 shows a slight delay in the viscosity increase.Figure 8: Viscosity of the used resinsThe infiltration behavior of the used resins is shown in Figure 9. Therefore, the droplet height is plotted against the interaction time. The broken lines represents the minimum of maximum value of the mearsurements, which was calculated on basis of at least 10 droplets. For a temperature of 20 °C a total infiltration of the droplet into the substrate cannot be detected for both resins.

For a temperature of 20 °C a total infiltration of the droplet into the substrate cannot be detected for both resins. The initial droplet height is higher for the resin with the trademark Araldite GY 764. This fact in combination with equal contact angles in Figure 10 lead to the assumption, that the volume of one drop of the resin GY 764 is higher than the volume of the droplet GY793. The droplet height decrease during the first 5 to 10 seconds faster than during the remaining time. With increasing temperature, the infiltration time decreased dramatically for both systems. For a temperature of 60 °C the total infiltration takes only approximately 2 seconds. This334acceleration of the infiltration process goes along with the reduced viscosity and thus the higher volume of liquid material, which can infiltrated the porous structure according to equation 2.

For layer times of approximately 40 seconds in selective laser sintering an infiltration time of only 2 seconds is favored because the infiltration itself must not enlarge the processing time.Figure 9: Droplet height dependent on time for different temperatures and resins, bulk density 0.39 g/cm³The infiltration behavior of the resins in a substrate with a higher porosity compared to Figure 9 is represented in Figure 10. The bulk density of the substrate changes from 0.39 g/cm³ to 0.36 g/cm³. This increase of porosity lead to an increase of infiltration speed, Figure 8 and Figure 9. At room temperature, the resin with the trademark Araldite GY764 infiltrated the porous structure in less than 60 second (Figure 10), whereas for a substrate bulk density of 0.39 g/cm³ after 80 seconds there is still a droplet height of 0.5 mm measurable.

For the substrate with the bulk density of 0.36 g/cm³ the deviation between the minimum and maximum values are larger than for the powder with a bulk density of 0.39 g/cm³. This may be a result of inhomogeneous packing and thus an irregular infiltration. The comparison of the two resins demonstrates a slightly faster infiltration in the case of Araldite GY 793. As shown before the infiltration speed increases with rising temperature due to the reduced viscosity of the resins.335For temperatures above 60 °C the absorption of the liquid is even faster and the contact angle as well as the droplet height cannot be detected any more.Figure 10: Droplet height dependent on time for different temperatures and resins, powder: PEBA bulk density 0.36 g/cm³The contact angle between the liquid resin and the bulk material is illustrated in Figure 11 for different resins, bulk densities and temperatures.

Compared to the droplet height the initial contact angle is at the same level for both resins. The contact angle decreases faster with increasing temperature for both substrates and resins. This can be traced back to the reduced viscosity in Figure 8. Furthermore, the contact angle shows almost the same tendencies like droplet height. For a temperature of 60° the contact angle tend to zero after less than 5 seconds and thus the process speed is not influenced dramatically in selective laser sintering.336Figure 11: Contact angle dependent on time for different temperatures, resins and bulk densitiesFigure 12 shows microscopy images of a cross section of a sample of PEBA with the resin GY 764. The samples were prepared in a selective laser sintering system type DTM Sinterstation 2000. The building chamber temperature was 130 °C (see DSC diagram Figure 4) and the liquid droplet and the powder material was exposed with a CO2 laser.

The building chamber temperature was 130 °C (see DSC diagram Figure 4) and the liquid droplet and the powder material was exposed with a CO2 laser. The images show, that the infiltration depth of the liquid is higher than the laser material interaction zone, which is visible as white areas. The used resin is absorbed into a thickness of approximately 1,000 µm. In regions far away from the surface the PEBA is not molten and the particles glue together due to the surrounding epoxy material. Nevertheless, these experiments show a principle feasibility of the represented hybrid process.Figure 12: Microscopy image of PEBA sample, which was infiltrated with a droplet of the resin GY 764337ConclusionsThis paper introduces a new hybrid additive manufacturing techniques, which will overcome the hurdle to manufacture multi material part in one additive process. With selective laser sintering of plastic only one material can be processed up to now.

With selective laser sintering of plastic only one material can be processed up to now. The implementation of reactive liquids, like epoxy resins, in the laser sintering process allow the generation of multi material parts. Therefore, the curing and infiltration behavior of the used resins has to be understood.The interaction of the resin with the porous powder bed is investigated with new test set up. The development of droplet height and the angle between the powder bed and the liquid were determined and linked to the infiltration behavior. For both analyzed epoxy resins Araldite 764 and Araldite 793 the Hardener XB 3473 was used. With increasing temperature the infiltrations process is accelerated for both resins. This goes along with the reduced viscosity with increased temperature. At 20 °C the resin does not infiltrate the powder bed in total. For the highest analyzed temperature of 60 °C the infiltration takes less than 2 seconds.

Assembly Automation. 201, pp. 39-46.[3] Schmidt J., et al.: Processing Behavior Of Spherical PBT Powders Produced By ANovel Process During Selective Laser Beam Melting. PPS 31, Jeju Island, Korea 2015.[4] Schmidt J., et al.: A novel process for production of spherical PBT powders and their processing behavior during laser beam melting. AIP Conference Proceedings. 2016; 140008.[5] Wegner A., et al.: Determination of Optimal Processing Conditions for the Productionof Polyamide 11 Parts using the Laser Sintering Process. International Journal of Recent Contributions from Engineering, Science & IT (iJES). 2015, pp.-5-12.[6] Amado A., et al.: Characterization and modeling of non-isothermal crystallization ofPolyamide 12 and co-Polypropylene during the SLS process. 5th International Polymers & Moulds Innovations Conference 2012. p. 207-16.[7] Verbelen L., et al.: Characterization of polyamide powders for determination of lasersintering processability.

The three-dimensional point clouds were obtained from outside of the build chamber with a fringe projection measurement system which was provided with access through an observation window. The curvature-based thresholding of powder bed point clouds demonstrates the ability to separate consolidated areas from the powder bed effectively. This segmentation of the point clouds with masks enables the detection of changes in the outline of consolidated areas between layers, computation of average drop due to the consolidation of the powder bed and separate analysis of both powder bed and consolidated areas. The high-level insights extracted from the point clouds could improve process control strategies, such as in-line defect detection during an additive manufacturing build as well as an in-process feedback system for tuning the optimal values of additive process parameters.

In summary, we show curvature-based thresholding as an effective segmentation for fringe projection point clouds, which can be further applied to detect defects, such as geometric defects and dimensional inaccuracy.the analysis ofKeywords: Additive Manufacturing, Polymer Powder Bed Fusion, In-Process Monitoring, Fringe Projection, 3D Point Cloud Processing, Curvature, SegmentationCorresponding author: aclaw@vt.edu1. IntroductionAdditive manufacturing (AM) has the potential to produce complex, functional components in small runs at unit costs competitive or better than other forms of manufacturing. To utilize these potential benefits, however, more research and development is essential in all aspects of production, ranging from a part material properties to its dimensional accuracy [1] and surface topography [2].

The specific AM process studied in this paper is polymer laser powder bed fusion (LPBF), which utilizes a laser beam as a thermal energy source to selectively fuse the areas of a polymer powder212bed. The bed is then lowered, and more powder is spread to repeat the process and build the parts in a layer-by-layer fashion.For industry to adopt LPBF processes for production, robust quality assurance must be established [3]. In recent literature, the framework of in-process monitoring and control is highlighted as a potential approach to address quality issues. Grasso and Colosimo emphasized the two crucial needs of automated defect detection and process control strategies for developing smart LPBF systems [4]. Everton et al. remarked there is industry demand for in-situ inspection and closed-loop control techniques for AM. They proposed that a priori knowledge of the part and build process is necessary for closed-loop control [5]. Mani et al.

They proposed that a priori knowledge of the part and build process is necessary for closed-loop control [5]. Mani et al. demonstrated the need for real-time control of powder bed fusion processes by understanding the correlations between process parameters, process signatures and product qualities [6].Given the significance of in-process monitoring to improve LPBF processes, the causes of common defects observed in LPBF processes have been reviewed. Malekipour and El-Monayri summarized typical defects existing in the LPBF process into four major categories, which are related to geometry and dimensions, surface quality, microstructure, and poor mechanical properties [7]. Grasso and Colosimo found six categories of defects, including geometric defects and surface defects, with four specified sources of defects for equipment, process, design for additive choices, and powder [4]. Similarly, Everton et al.

Similarly, Everton et al. listed five major LPBF discontinuities which are porosity within the material, as well as between layers, balling, unfused powder, and cracking [5].A significant challenge of in-process monitoring associated with the LPBF processes is the difficulty to characterize part integrity using conventional methods, such as with a coordinate measuring machine (CMM), because LPBF processes typically take place within a closed chamber with limited space, and contact measurements could interfere with the process. Fringe projection, a non-contact and three dimensional (3D) measurement technique, has been proposed for surface metrology of metal AM parts [2] [8] as well as being implemented as an in-situ powder bed measurement technique [9]. However, there is limited reporting of deploying fringe projection as part of an in-process monitoring system. Zhang et al.

However, there is limited reporting of deploying fringe projection as part of an in-process monitoring system. Zhang et al. proposed the implementation of fringe projection to measure the surface topography of a metal powder bed and showed that height maps contain information about the fusion process [9]. In addition to capturing the 3D surface of the powder bed area, Li et al. demonstrated that polymer powder beds are amenable to fringe projection system as contours of consolidated areas can be calculated for in-process control algorithms [10].Unlike many 3D measurement techniques that only use cameras, fringe projection produces quantified measurements of the entire powder bed with minimal pixel correspondence issues. The data captured contain a representation of the surface in the form of a 3D point cloud. The information collected is helpful to derive insights about consolidation defects and dimensional inaccuracy.

The information collected is helpful to derive insights about consolidation defects and dimensional inaccuracy. However, there is no perfect sensing system which can capture all types of defects. Hence, the fringe projection technique has its limitations, and one of these is the time taken to capture a large set of data points, which213might be longer than the typical layer build time. A typical scan of the fringe projection technique can take approximately 10 s, while the time between the beginning of one layer and the beginning of the next layer in the studied system is at least 10 s and becomes longer with larger and more complex parts. The time taken for the analysis of point clouds is needed to keep within the layer build time for understanding the current layer condition before proceeding to the next layer. However, this limitation can be solved by downsampling the required points needed for timely analysis of point clouds.

However, this limitation can be solved by downsampling the required points needed for timely analysis of point clouds. Apart from the disadvantage of high acquisition time, fringe projection only measures the surface of the powder bed which has little to no information of the layers underneath [11].Another current limitation of fringe projection as an in-process monitoring system is its need for calibration, which is required to map the measured points into real-world coordinates [12]. As well as a potentially challenging system calibration, availability of point cloud analysis techniques explicitly applied to the application of in-process monitoring is lacking. Additionally, the data structures and processing methods of 3D point cloud data are different from those applied to 2D imaging systems [13]. In a recent study, Southon et al.

Additionally, the data structures and processing methods of 3D point cloud data are different from those applied to 2D imaging systems [13]. In a recent study, Southon et al. measured maximum height deviation from the powder bed measurement point clouds to demonstrate the capability of the fringe projection system to detect defects early, such as random defects and curling [14]. Southon et al. highlighted the potential of utilizing point clouds for process signature extraction as part of in-process monitoring and control system.This work aims to demonstrate that curvature-based analysis of point clouds canproduce useful features for detecting defects related to geometry and surface quality.2. MethodsThe term point cloud is defined as a set of points represented in a coordinate space. Rusu and Cousins explained the term as a data structure which represents a collection of multidimensional points [15].

Rusu and Cousins explained the term as a data structure which represents a collection of multidimensional points [15]. Weinmann further defined the term as a collection of 3D points, which are characterized by spatial coordinates which may be assigned with additional attributes, such as intensity, color, thermal, and other information [16]. The point clouds in this study are acquired by a fringe projection system that projects structured light onto the part’s surface and records the distortions of the light pattern induced by the part’s geometry (the system is further detailed in [11]). The distorted patterns are captured by a camera and used in the reconstruction of the surface geometry using the least-squares phase shifting algorithm [17].The topic of curvature estimation of 3D surfaces represented as point clouds can be found in literature as early as Kreyszig in 1959, who discussed a detailed overview of concepts related to the differential geometry of 3D surfaces [18].

Local primitives, such as surface normals, curvature, and the change of curvature may provide useful information for the understanding of scanned parts in the form of a point cloud. Both Porteous [19] and Giblin [20] studied the use of curvature information for analyzing and studying surface shape. Douros and Buxton presented a review on the curvature estimation topic and proposed a new surface curvature estimation using a quadric surface representation [21].214Bae and Lichti proposed using the change of curvature and approximate normal vector of the surface to determine the registration of point clouds [22]. Jovančević et al. presented a defect detection and characterization system based on local normal and curvature properties of point clouds for airplane exteriors [23]. Madrigal et al. proposed a model-based 3D descriptor based on surface normals of point clouds for automatic surface inspection [12].

Madrigal et al. proposed a model-based 3D descriptor based on surface normals of point clouds for automatic surface inspection [12]. Both articles approach the processed point clouds as 2D images which can be used for defect detection and surface inspection. Furthermore, Bartkowiak et al. studied the use of multi-scale areal curvature analysis on the surfaces fabricated by fused deposition modeling (FDM), one of seven types of AM processes identified by ISO/ASTM52900-15 [24].The first feature used for the segmentation of point clouds in this paper is curvature 36T), where 36Tare the maximum and the minimum curvatures of the surface at a point [25]. Gaussian curvature K is defined as . 𝛋𝛋1 For estimating the curvature of a surface in 3D space, the quadratic approximation of a ), assuming f is , surface S [18] is used in the neighborhood of a point P( ).

𝛋𝛋1 For estimating the curvature of a surface in 3D space, the quadratic approximation of a ), assuming f is , surface S [18] is used in the neighborhood of a point P( ). The equations for the Gaussian curvature and the mean twice differentiated at ( curvature of S at P are defined as follows [26]information. Curvature characterizes a point by its principal curvatures (36T,36Tand𝛋𝛋2 36T, while mean curvature H is defined as𝛋𝛋1𝛋𝛋2𝛋𝛋1𝛋𝛋2),,𝑥𝑥0𝑦𝑦0𝑓𝑓(𝑥𝑥0, 𝑦𝑦0𝑥𝑥0𝑦𝑦0𝐾𝐾 = 𝛋𝛋1𝛋𝛋2 =𝑓𝑓𝑥𝑥𝑥𝑥 𝑓𝑓𝑦𝑦𝑦𝑦− 𝑓𝑓 2 2(1+𝑓𝑓𝑥𝑥+𝑓𝑓2𝑥𝑥𝑦𝑦 2𝑦𝑦),(1). (2)22𝑥𝑥�𝑓𝑓𝑦𝑦𝑦𝑦−2𝑓𝑓𝑥𝑥 𝑓𝑓𝑦𝑦𝑓𝑓𝑥𝑥𝑦𝑦+(1+𝑓𝑓𝑦𝑦)𝑓𝑓𝑥𝑥𝑥𝑥�1 +𝑓𝑓𝛋𝛋1+ 𝛋𝛋23 22 = Where f and its subscripts are used to indicate the first and second order local derivatives in x and y. The second feature used in this work is prominence variation.

The second feature used in this work is prominence variation. This is calculated as the distance between each point and a locally fitted, least-squares mean plane [27], given by all neighboring points within a sphere of kernel radius.𝐻𝐻 =222(1+𝑓𝑓𝑥𝑥+𝑓𝑓𝑦𝑦)The algorithm flow to calculate curvature and prominence variation is as follows: for each point P in the point cloud, a neighborhood of points, within the kernel radius, is selected. The kernel radius is defined as the radius of a sphere within a neighborhood of points for the curvature estimation and prominence variation calculation. With the points selected within the kernel radius, the fitted quadratic surface is computed and, therefore, the curvature values for the central point are calculated. For the prominence variation calculation, a fitted plane based on the least-squares method is computed for each point using the neighbors within a chosen kernel radius.

For the prominence variation calculation, a fitted plane based on the least-squares method is computed for each point using the neighbors within a chosen kernel radius. The prominence variation values of each point are calculated based on the distance between each point and this plane. Last, the curvature values and prominence variation values are assigned back to the point P as additional information.215𝛋𝛋𝟏𝟏 + 𝛋𝛋𝟐𝟐23. Experimental Study3.1 Fringe projection data acquisition systemThe fringe projection system used was a commercial white light scanner, NUB3D SIDIO XR. The nominal measurement pitch was 75 μm in all three orthogonal directions, with a build chamber volume of 200 mm × 150 mm × 90 mm, and a stated volumetric accuracy after calibration of 15 μm. The fringe projection system acquired the measurements through the operator observation window of a EOS P100 polymer powder bed fusion machine.

The fringe projection system acquired the measurements through the operator observation window of a EOS P100 polymer powder bed fusion machine. For the in-process measurements of the powder bed surface, the viewing angle was roughly 55° from horizontal. The builds consisted of single tensile test specimens, with best measurement results being when the test specimens were closer to the front of the build chamber, due to reduced data drop out in the sections furthest from the fringe projection system. The build of a single tensile test specimen took thirty layers to finish and then the 3D scanning process continued for at least ten layers.3.2 Point cloud curvature analysis flowThe point cloud measurements captured the powder consolidation due to sintering after each laser beam passed and the curvature values and prominence variation values of the point cloud were processed.

Next, the point cloud measurements with the added information of curvature and prominence variation were converted into 2D images using the scattered interpolant method [28]. The 2D images showed the curvature maps and prominence variation map of both part and powder bed. At this stage, the curvature maps and the prominence variation map may highlight defects that deviate in term of shape, dimension or any unwanted surface geometry. The curvature maps were further processed for mask creation. The region of interest was specified as only the parts for the reduction of point clouds. The cropped regions from curvature maps and prominence variation map were then binarized based on Otsu’s thresholding [29], to obtain the part outline. A final mask representing the outline of a part was obtained by dilation and erosion.The purpose of the mask obtained in the previous step is to segment the parts from the powder bed area for defect analysis.

A successful segmentation enables the analysis of any geometric defects occurring either in the parts or powder bed during each layer. The mask resulting from segmentation was applied to the height map to calculate average drop due to consolidation for the part. Furthermore, a different mask can be produced for each layer. From this, the mask of each layer was subtracted from the mask of the previous layer to detect the outline change over layers. In addition, a separate analysis of powder bed and the part can be undertaken for defect detection and defect characterization.2164. Results and Discussion4.1 Curvature maps and prominence variation mapThe curvature-based analysis methods are implemented for the acquired powder bed point clouds. Each 3D scanned layer of the single tensile strength specimens contains approximately 290 000 points.

Each 3D scanned layer of the single tensile strength specimens contains approximately 290 000 points. Both the curvature maps and the prominence variation map, which reveal catastrophic defects, warping, and geometric defects, are presented below.4.1.1 Catastrophic defectsFigure 1 shows three surface analysis maps of the sintered layer of a part with observed debris on the right-hand side. The debris is a defect which could affect the part negatively. The debris is detected on the 16th layer of the print, where the right-hand side of the part is likely disrupted by the powder spreading blade. The unwanted debris is identified in all three surface analysis maps.(a)(b)(c)Figure 1. (a) Gaussian curvature map, (b) mean curvature map, c) prominence variation map of the 16th layer of a failed build with debris discovered on the right-hand side of the part.

(a) Gaussian curvature map, (b) mean curvature map, c) prominence variation map of the 16th layer of a failed build with debris discovered on the right-hand side of the part. This tensile test piece was printed at 165 °C, 17 W laser power with a laser scan speed of 2500 mm/s and a hatch spacing of 250 µm.2174.1.2 WarpingFigure 2 (a), (b) and (c) show three surface analysis maps of the final sintered layer of a part, built with a laser power of 4 W, with severe warping on both ends. For the studied part, warping, another form of defect which is common in the LPBF process [30], is observed on many layers including the 30th layer of the specimen, see figure 2. Furthermore, heaping of surrounding powder can also be identified at about -30 mm in the x-direction. This phenomenon is similar to the observation made by Southon et al. in [14]. In an ideal LPBF process, the part should lay flat on the powder bed after laser sintering.

This phenomenon is similar to the observation made by Southon et al. in [14]. In an ideal LPBF process, the part should lay flat on the powder bed after laser sintering. However, the curling effect causes both ends of the part to be elevated. Therefore, the curvature of the warping issue is easily identified in all three surface analysis maps.Even after the part has finished printing, the severity of warping can still be observed throughout the subsequent layers. Figure 2 (d), (e) and (f) show that the continuing effect of the curled ends of the part remains even after ten layers of new powder have been spread on top of the part. The elevated edges of the parts are observed in all three analysis maps while the center portion of the part is buried underneath the powder. This observation suggests that curvature-based analysis is effective at identifying features related to the issue of warping.(a)(d)(b)(e)(c)(f)Figure 2.

This observation suggests that curvature-based analysis is effective at identifying features related to the issue of warping.(a)(d)(b)(e)(c)(f)Figure 2. (a) Gaussian curvature map, (b) mean curvature map, (c) prominence variation map of the final sintered layer (30th layer) of a successful build with warping on both ends of the part. (d) Gaussian curvature map, (e) mean curvature map, (f) prominence variation map of the 40th layer of a build with the continuing effect of warping on both ends of the part.2184.1.3 Geometric defects in the partFigure 3 (a), (b) and (c) show three surface analysis maps of the sintered layer of a part with a surface protrusion on the left edge of the top part at the 19th and 28th layers. The surface protrusion, which appears to be a geometric defect, is observed to begin on the 19th layer of the print and continually developed into a larger protrusion over subsequent layers.

The surface protrusion, which appears to be a geometric defect, is observed to begin on the 19th layer of the print and continually developed into a larger protrusion over subsequent layers. The prominence variation map seems to be less sensitive to surface defects, where only a smaller region of dots indicates the variation on the local surface. It should also be noted that a laser scanning line artifact, shown in Figure 3 (c) and (f), is captured on the prominence variation map. This laser scanning line artifact is likely to be a combination of individual laser scan lines because about 68 lines can be seen across the 165.2 mm length of the part, leading to a spacing of 2429.4 µm which is significantly larger than the 250 µm scan spacing of the laser path.In the parts measured which were printed at 165 °C with higher laser power, 27 W and the same scan speed of 2500 mm/s, a random defect occurred on at the 19th layer.

The observation suggests that a geometric defect, such as a surface protrusion, could occur randomly even with process parameters close to optimal. Hence, the fringe projection can satisfy the need for in-process monitoring to detect random defects.(a)(d)(b)(e)(c)(f)Figure 3. (a) Gaussian curvature map, (b) mean curvature map, (c) prominence variation map of the 19th layer where a surface protrusion, as marked with a red circle, on the left side of the top part. (d) Gaussian curvature map, (e) mean curvature map, (f) prominence variation map of the 28th layer where the observed surface protrusion is larger and in the same region of the top part.2194.2 Curvature based segmentation for mask creationThe curvature and prominence variation maps are useful for defect identification on each acquired layer point cloud. However, the need to perform segmentation of parts from the powder bed is critical in differentiating the location of defects occurring in each layer.

However, the need to perform segmentation of parts from the powder bed is critical in differentiating the location of defects occurring in each layer. Therefore, the successful segmentation is essential to recognize the severity of defects by their location. The following results present the curvature-based segmentation with a mask created using Gaussian curvature, mean curvature and prominence variation, which captures the shape of the part. Figure 4 (a), (b) and (c) show the results of the three initial masks after binarizing the respective maps based on Otsu’s thresholding.In figure 4 (a), the mean curvature-based mask is successful at detecting the edges of the parts; the black region between two white regions indicates the edge of the part, while white regions indicate areas surrounding the edge.

In figure 4 (b), the Gaussian curvature-based mask consists of several blobs surrounding the edge of the part, and the blobs are useful for corner detection, where the corner of the mean curvature-based mask is missing. In figure 4 (c), the prominence variation-based mask performs better than the other two masks at outlining the edge of part despite noisy points.In summary, each initial mask provides additional cues, and they complement each other when creating the final mask. With the combination of the three initial masks, the final mask can capture the complete shape of a part, which is shown in figure 4 (d) and (e).(a)(d)(b)(e)(c)Figure 4. (a) mean curvature mask (b) Gaussian curvature mask (c) prominence variation mask after binarizing surface analysis maps based on Otsu’s thresholding.

(a) mean curvature mask (b) Gaussian curvature mask (c) prominence variation mask after binarizing surface analysis maps based on Otsu’s thresholding. (d) the result of a combination of three initial masks which represent the shape of part (e) the final mask is created after implementing the image processing methods of erosion and dilation.2204.2.1 Computation of drop due to consolidationBy applying the curvature-based mask to segment the part from the powder bed, the analysis of powder bed point clouds for an in-process monitoring system can be improved. One of the process signatures for in-process monitoring is drop due to consolidation, which provides a basic understanding of the quality of consolidation of layers [9]. Figure 5 and figure 6 show the segmentation mask applied to the height map and the computation of the average drop due to consolidation in each layer for point clouds taken at a laser power of 17 W. As Zhang et al.

As Zhang et al. discussed, the average height drop is expected to converge to a constant value after an initial decrease in the first few layers [1].Figure 5. Example of a successful segmentation masking applied to a powder bed height map to compute the average drop due to consolidation. However, there is a small portion of the powder bed captured within the mask. The information of powder bed surrounding the part might be accounted into the computation of the average drop which leads to A slight inaccuracy of the computation. However, this issue can be solved by filtering out the positive values which appear to be part of the powder bed region.Figure 6. Average drop due to consolidation for the test part printed at 165 °C with a laser power of 17 W and a scan speed of 2500 mm/s. Drop due to consolidation increases with increasing layer count before layer 16 where the debris, shown previously, occurs.

Drop due to consolidation increases with increasing layer count before layer 16 where the debris, shown previously, occurs. From the 16th layer to the 30th layer, the average drop due to consolidation showed a similar trend. From the 31st layer to the 40th layer, the positive values of drop indicate that the part is done and a new layer of powder is deposited, without laser sintering, on each layer.2214.2.2 Binarized outline change over layersThe powder bed of each layer can be separated into part and powder bed using the segmented mask. By subtracting each layer’s segmentation mask from the previous layer mask, a three-valued map, showing the shifts of the contour defining the boundary of the parts, is created. The two signed values (-1, +1) indicate as the change of the previous layer and the current layer while 0 value indicates no change between layers. This technique can help identify powder heaping, which is likely caused by ineffective bed-spreader contact dynamics [31].

This technique can help identify powder heaping, which is likely caused by ineffective bed-spreader contact dynamics [31]. Figure 7 shows the detection of alternating heaping of powder from the 13th layer to the 15th layer as well as the detection of the deviation of part mask outlines between layers.Figure 7. From the 13th layer to the 15th layer, fringe projection point clouds capture the alternate pattern of heaping of powder. The EOS P100 uses a straight metal powder spreading blade that travels in an arc across the powder bed. The accumulated powder at the alternating ends suggests that the powder spreading blade might not wholly smooth the powder over the submerged, consolidated part. From the 16th layer to the 18th layer, a series of layer changes is shown with a noticeable deviation from the outline shape of the part.4.2.3 Separate analysis of powder bed and partAnother contribution of the segmentation mask is to enable separate analysis of the powder bed and part.

Separate analysis of powder bed and part is significant in improving the defect classification. Figure 8 (a) and (b) show the powder bed and the part at layer 16, where the debris is detected within the part and also in the powder bed. Figure 8 (c) and (d) show the curling effect is observed on the powder bed and the part on the 31st layer. The222ability to split out the part and the powder bed is critical because the separate analysis can show the location of defects. By knowing the location of defects, the severity of the defects can be determined and can inform the system whether to continue or stop the printing process.(a)(c)(b)(d)Figure 8. (a), (b) Separate analysis of the powder bed and the part where debris is observed on the 16th layer.

(a), (b) Separate analysis of the powder bed and the part where debris is observed on the 16th layer. (c), (d) Separate analysis of powder bed and the part where warping is observed on the 31st layer.ConclusionIn summary, this work demonstrates that the curvature analysis of point clouds acquired by a fringe projection system can be used for in-process monitoring of powder bed fusion to identify catastrophic defects, warping, and geometric defects. Furthermore, the curvature and prominence variation information can be used to create segmentation masks for the powder bed and part. This study has demonstrated that a segmentation mask with no a priori part information can be created and can be used to compute drop due to consolidation in the region of interest. Also, the segmentation masks enable detection of changes in the outline of the consolidated areas between layers by comparing the outline of successive layers.

Also, the segmentation masks enable detection of changes in the outline of the consolidated areas between layers by comparing the outline of successive layers. The alternate pattern of heaping of powder and outline shape deviation of a layer can also be observed. Another contribution of curvature-based segmentation is the separate analysis of the powder bed and the part which allows for localization of defects.The insights extracted from the point clouds analysis can be used to inform process control strategies. The drop due to consolidation is a significant process signature for determining the quality of layer processing [14]. Being able to compute the average drop from the height map with an automatic segmentation mask, provides information that can be used to further the understanding of polymer LPBF. Also, the binarized outline change over layers could be utilized as an alert for early defect detection during a build.

Also, the binarized outline change over layers could be utilized as an alert for early defect detection during a build. The mask analysis of the outline change of each layer can be studied for feature extraction suitable for machine learning classification applications. The output of the separate analysis of both powder bed and the part can further identify defect types and defect characterization. With a priori knowledge of the potential causes of each defect type, an in-process closed-loop223feedback system could be implemented for tuning the optimal values of LPBF process parameters.The aim of the next implementation of this research is using curvature-based segmentation of LPBF point clouds as a real-time analysis tool for in-process defect detection and classification. A new study on defect classification and associated causes is ongoing.

A new study on defect classification and associated causes is ongoing. In conclusion, the application of the novel framework developed herein using a fringe projection system for in-process monitoring and analysis in a powder bed fusion AM process will be performed on a real case study.AcknowledgmentsThis work was supported by the Engineering and Physical Sciences ResearchCouncil [grant number EP/M008983/1, EP/L01534X/1]References[1] W.S. Land, B. Zhang, J. Ziegert, A. Davies, In-situ metrology system for laser powder bed fusion additive process, Procedia Manuf. 1 (2015) 393–403. doi:10.1016/j.promfg.2015.09.047. [2] multi-sensor in-situ inspection system for additive manufacturing, Proc. ASPE/euspen Advancing Precision in Additive Manufacturing. (2018) 248-242. [3] Charitidis, Additive manufacturing: scientific and technological challenges, market uptake and opportunities, Mater. Today. 21 (2018) 22–37. doi:10.1016/j.mattod.2017.07.001. [4] M. Grasso, B.M.

rved Layer LOM of Ceramics and CompositesDonald A. Klosterman, Richard P. Chartoff, Nora R. Osborne, George A. Graves, Allan Lightman, Gyoowan Han, Akos Bezeredi, Stan Rodrigues Rapid Prototype Development Laboratory University ofDayton, Dayton, OHSung Pak, Gary Kalmanovich, Leon Dodin, Song Tu Helisys, Inc., Torrance, CAABSTRACTA novel rapidprototyping (RP) technology incorporating a curved layer building style has been developed. The new process,based on Laminated Object Manufacturing (LOM), is suited for efficient fabrication of curved layer structures made from ceramics and fibrous composites. Anew LOM machine was developedithatuses ceramic tapes and fiberprepregs as a feedstock and outputs at11fee dimensional green form. The green ceramic is then processed to a seamless, fully dense ceramic structure using traditional ceramic techniques. This report summarizes the new LOM process. and necessary hardware.

The green ceramic is then processed to a seamless, fully dense ceramic structure using traditional ceramic techniques. This report summarizes the new LOM process. and necessary hardware. Also reviewed is the development of ceramic preforms and accompanying process technology for net shape fabrication of ceramic matrix composites (CMCs). Compared to making curved objeds with the standard flat .layer LOMprocess, the curved process affords the advantages of eliminated stair-step effect, increased build speed, reduced waste,reduced need for decubing, and the ability to maintain continuous fibers in the direction of curvature.CURVED LAYER LOMBased on LOM, a novel RP technology has been developed to provide an enhanced capability for fabricating structural ceramic and composite articles. A comprehensive research program was recently completed that involved materials and process research, machine design,and system software adaptation [1, 2].

A comprehensive research program was recently completed that involved materials and process research, machine design,and system software adaptation [1, 2]. The envelope of available RP materials was expanded to include structural composites. In addition, an entirely new building paradigm was implemented: thecurvedlayer building style. Instead of being limited to building with flat layers, the LOM machineis now capable of building in a curved-layer manner. The new curvedJayer LOM process allows continuous fiber composites to maintain fiber continuity in the plane of curvature in order to achieve optimum mechanical performance.The curved layer LOM process originated from the need to fabricate fiberreinforced structures containing. sloping, curved s\lffaces, especially thin curved-shell components. An important factor in these structures is maintaining fiber continuity inthe curved surfaces.

sloping, curved s\lffaces, especially thin curved-shell components. An important factor in these structures is maintaining fiber continuity inthe curved surfaces. All RPprocesses are capable of fabricating complex, curved geometries using thin flat layers in combination with post machining of the final part. However, flat671layer RP processes are incapable of addressing the larger geometrical issues involved with fiber composite fabrication, namely fiber orientation and continuity. In addition to the technical incentives for·implementing a curved layer build style. there is an economic incentive.. Such a process will have a favorabl~cost benefitpayoffbecauseofthe reduced raw material costs compared to flat layer processes· (see Figure 1). This results from a dramatic reduction in the amount of waste material generated during processing.The curved layer LOMprocess is illustrated in Figure 2. It begins withproduction of a matched tool or mandrel for the intended part.

It begins withproduction of a matched tool or mandrel for the intended part. The temperature and pressure requirementsfor this mandrel are not demanding, so it can be made with the standard flat layer LOMprocess usingLOM paper. The finished mandrel is mounted to the flat building platform in the curved layer LOM machine. Sheets of the desired build material are loaded onto a rotatable feed table, picked up with a vacuum chuck, and transported to the mandrel. A flexible thermoforming mechanism laminates each new layer to the curved mandrel with steady, uniform pressure. A laser cuts each layer accounting for the sloped surface. The fiber orientation can be varied from layer to layer by programming the rotatable feed table. The process proceeds one layer at a time until the part is finished. The part is then removed from the mandrel and the excess material is manually stripped away ("decubed").

The process proceeds one layer at a time until the part is finished. The part is then removed from the mandrel and the excess material is manually stripped away ("decubed"). The advantages ofthe curved layer process are elimination of stair step effect and improved surface quality, increased build speed, reduced waste, and easier decubing.Since the new process is a significant departure from the flat layer LOM process, substantial neW hardware and software developm~nt was necessary. The software issues are summarized in a previous report [3]. New hardware components inc1udedthe material sheet feeding and rotating mechanism, curved layer bonding apparatus, and curved surface laser cutting. An automatic sheet feeding system rather than a roll feed was desired for two main reasons. First, many commercially available advanced materials, such as prepregs (preimpregnated fiber preforms) and ceramic tapes, are stiff at room temperature and/or available only in sheet form.

First, many commercially available advanced materials, such as prepregs (preimpregnated fiber preforms) and ceramic tapes, are stiff at room temperature and/or available only in sheet form. Second, there is more material waste associated with the automatic roll feed system. Although this waste may be affordable when building LOM-paper parts, efficient use of material is critical when working with costly feedstocks such as fibrous composites and ceramics.The curved layer bonding mechanism is a flexible, resistively heated pad that is backed with a silicone rubber frame (Figure 3). The frame consists of a top and bottom horizontal surface connected by vertical walls. The placement of the internal walls was custom designed to provide even pressure to the curved layer parts under consideration in this study. Vacuum cups protrude vertically though the rubber frame in order to pick up a new material layer.

Vacuum cups protrude vertically though the rubber frame in order to pick up a new material layer. A single layer is held just underneath the heating surface as the entire assembly is moved over the mandrel via a rail system. The new layer is laminated by elevating the build platform up into th~ thermoformer and holding it in contact for about one minute. The bonding pressure is approximately 2 - 10 psi, which is adequate for laminating typical tapes and prepregs. Figure 4 shows a ceramic part that has been layed-up on a LOM paper tool.672Cutting curved layer parts required new computer algorithms for coordinatingsimultaneous control of the laser beam in the X, Y, and Z directions. In a conventional LOM machine, laser cutting is performed using a plotter system that transports a mirror and focusing lens over the X-Y envelope.

In a conventional LOM machine, laser cutting is performed using a plotter system that transports a mirror and focusing lens over the X-Y envelope. The distance between the focusing lens and the top of the part is adjusted only once after placing each new layer, ensuring that the laser focus is maintained on the horizontal part surface. With the curved layer LOM machine, the build platform must be translated up and down dynamically in order to maintain the laser focus on the curved surface. An eight axis motion control card and new software algorithms were necessary to control this motion smoothly and quickly. A 225-Watt Diamond (pulsed) C02 laser was installed on the machine and proved effective in cutting SiC fibers and ceramic tapes.

A 225-Watt Diamond (pulsed) C02 laser was installed on the machine and proved effective in cutting SiC fibers and ceramic tapes. The quality of the fiber cuts was not as good as that obtained with a copper vapor laser as previously reported [1,4], but the Diamond laser is more reliable, smaller, and less expensive than the copper vapor laser.CMC PREFORM DEVELOPMENT FOR LOMSiC/SiC was used as a focus "model" material system for demonstrating curved layer, net shape CMC fabrication. A novel approach was used to fabricate CMCs that involved lay-up of separate, alternating layers of ceramic tape and fiber prepreg [1]. Green parts containing this alternating layer architecture were made using LOM, subsequently pyrolyzed, and densified through reaction bonding.Fiber prepregs were fabricated with unidirectional SiC fiber tows (Nicalon™,Dow Corning, carbon coated grade) and a furfural-phenolic thermosetting resin (FurCarb® UP440, QO Chemicals, Inc).

The furfural resin served a dual role: as a binder and adhesive during the part fabrication and as a carbon source during the subsequent reaction bonding process. Descriptions of the reaction bonding process and monolithic SiC tapes are given in a previous report [1]. Alternating layers of the fiber prepreg and ceramic tape were delivered to the LOM machine as one single sheet, referred to asa "tape-preg" [1]. Tape-pregs, made by pre-adhering aprepreg layer to a tape layer, were stiff, non-tacky, and board like at room temperature, becoming soft and pliable at higher temperatures (130°C). The tape-preg contained approximately 25% fibers by weight, and overall it proved to be a robust sheet material suitable for the LOM process.The advantages of the alternating layer technique are the relative ease ofpreparation of the preforms, elimination of fiber abrasion from the ceramic particles, reasonably high fiber volume fraction in the final CMC, and low cost resin binder.

The tape-preg embodies several important characteristics of other commercially available CMC preforms. Key among these characteristics is the use of a fiber prepreg that contains a thermosetting, ceramic precursor resin. Based on this similarity, it is anticipated that LOM will accommodate other commercially available CMC preforms as easily as the "tape-preg".Several parts were manufactured with curved layer LOM using SiC/SiC tapepregs. Acceptable process parameters (e.g., temperature, pressure, time) were673determined through initial experimentation with a manual operation,. mock~up system at the University of Dayton [4], and finally through operation of the actual curved layer LOM machine. A bonding time of 1· minute at 130~150°C. was adequate for the tape~ pregs;thebondingpressure was estimatedto be 5 psi. Several curved layer parts were produced and successfully decubed. These green composites were not fully cured, but theywere stiffandself-supporting.

Several curved layer parts were produced and successfully decubed. These green composites were not fully cured, but theywere stiffandself-supporting. Curved layer LOMwas also used to fabricate lTIonolithic SiC parts by laminating only the ceramic tapes (no fibers}. The temperature in this case was 90°C, and butanol was sprayed between the layers to act as a tackifier. Thus, the LOM process was successfully used to produce various types of green ceramic forms that were robust and handleable. Additional processing steps were required to convert them to full density.POST-LOM PROCESSINGIt is not possible to execute all ofthe necessary ceramic processing steps within theLOM machine. The LOM machine is used only to produce the green form. Thus, post processing steps are needed. The overall process used for SiC/SiC is illustrated in Figure 5.The first post-processing step combines binder burnout, pyrolysis, and pressinginto a single operation.

The overall process used for SiC/SiC is illustrated in Figure 5.The first post-processing step combines binder burnout, pyrolysis, and pressinginto a single operation. Although green LOM parts are sufficiently laminated to survive LOM processing, decubing, and handling, they require additional layer consolidation prior to the final densification step. Pressure must be applied during binder burnout and pyrolysis to counteract delamination and bloating that result from several sources such as outgassing and relaxation of residualstresses imparted during the lamination step. In order to maintain the complex geometry of green parts during pressure application, a technique involving quasi-isostatic powder pressing was developed.Green parts from the LOM process were placed in a cylindrical chamber whichwas subsequently back filled with powder and fitted with a ram. Monolithic SiC was used as the powder medium in the chamber.

Monolithic SiC was used as the powder medium in the chamber. For LOM parts that could be fitted with a porous Teflon bag, silica was used as the powder medium (Le., sand pressing). A heated, programmable, uniaxial press was used to apply a heated pressure cycle to the enclosed chamber. The powder medium enabled pressure to be evenly distributed while allowing volatile degradation products to escape through a permeable bed. The various SiC/SiC and monolithic SiC parts were pressed at 60 psi and 30 psi, respectively, and at temperatures up to 325°C. Shrinkage of SiC/SiC and monolithic SiC parts during this step was 7% ±2% and 5% ±20/0 respectively, in the z direction (parallel to press movement), and zero in other directions.

Shrinkage of SiC/SiC and monolithic SiC parts during this step was 7% ±2% and 5% ±20/0 respectively, in the z direction (parallel to press movement), and zero in other directions. Appropriate scaling of the computer graphic file for part building in the z direction by 7% or 5% will compensate for this dimensional change due to compression.An additional, freestanding binder burnout and pyrolysis step up to 700°C inargon was needed to achieve full chemical conversion of the organic resins to carbon. Because most of the total expected binder burnout and resin pyrolysis occurred during the powder pressing step, no part damage or shrinkage accrued during freestanding burnout.674The resulting porous structures were densified through reaction bonding. The final parts are shown in Figure 6.The microstructure of flat LaM SiC/SiC panels prior to pyrolysis illustrates thealternating tape/fiber layer arrangement, excellent compaction, and minimal porosity (Figure 7, left).

The photomicrograph of a specimen that has undergone reaction bonding indicates that there is some porosity, although much of this may be attributable to particle and fiber pull-out that occurred when the specimen was polished. It is clear, however, that the flatness, continuity, and integrity of the layers have been well maintained. The level of porosity is comparable to that of commercially available CMC systems.Upon closer examination of the LaM SiC/SiC microstructure, it is evident thatalthough the fibers are intact, there is fiber damage particularly at the fiber interface (Figure 8). This damage occurs as a result of the instability ofNicalon fibers at the temperatures encountered in the reaction bonding process (up to 1600°C).

This damage occurs as a result of the instability ofNicalon fibers at the temperatures encountered in the reaction bonding process (up to 1600°C). There are several potential solutions to this problem, such as developing superior fiber coatings and using silicon alloys that will infiltrate the composite at lower temperatures..However, the physical mating of the layers is intimate and the overall infiltration efficiency is quite good relative to commercial CMC systems.Testing of final SiC/SiC specimens in four-point bend was not possible due todamage to specimens during cutting. The brittle behavior of the samples was expected, since the fiber coating problem has not been solved. The major result to report here is that the entire process, starting with fiber preforms and ending with a near net shape, densified part, has been fully demonstrated.

The major result to report here is that the entire process, starting with fiber preforms and ending with a near net shape, densified part, has been fully demonstrated. Densified samples such as the flame holder and body armor are handleable, and photomicrographs illustrate that the microstructure compares quite favorably to commercial CMC systems. Further work with other material systems is being carried out to fully characterize and assess the overall process reproducibility and accuracy.SUMMARY AND CONCLUSIONSThe feasibility of using LaM for net shape, freeform fabrication of monolithicceramics, continuous fiber CMCs, and curved layer composites has been fully demonstrated. Monolithic SiC and continuous fiber SiC/SiC were used as demonstration material systems. Commonly available preforms such as ceramic tapes and fiber prepregs are entirely suitable as feedstocks to the process.

Commonly available preforms such as ceramic tapes and fiber prepregs are entirely suitable as feedstocks to the process. The overall process methodology involves use of a LaM machine to produce dimensionally accurate green forms directly from CAD files, followed by post processing steps to bring the part to fulldensity. Both flat layer and curved layer laminates can be built with this technology. The curved layer paradigm is critically important for fabricating continuous fiber, curved shell composites in order to maintain fiber continuity. Curved layer SiC/SiC composites were produced with good microstructure, although mechanical properties. are n.otexpected to be as high as commercial CMC systems due to ineffective fiber coatings. Inherent to the process of making the CMCs was the production of a precursor composite structure, consisting675essentially of a polymer matrix composite (PMC), which was subsequently converted to a CMC.

Thus, the results demonstrate that PMC fabrication is also viable with the new LOM process. This also was demonstrated in our previous LOM processing studies using epoxy resin prepregs [5].Although the parameters of the various processing steps are expected to change withdifferent materials systems, the overall LOM process is expected to be generically applicable to a wide range of high performance materials. Thus, an entirely new fabrication capability is now available for designers and manufacturers. The new LOM process holds promise for lowering the cost of fabricating monolithic ceramic, CMC, and polymer matrix composite parts by virtue of its automation, which reduces fabrication time and eliminates the need for tooling. The process can best be applied by industry as a product development tool for fabricating molds, tools, or testable prototypes, or used directly for small lot production.This is an ongoing effort.

The process can best be applied by industry as a product development tool for fabricating molds, tools, or testable prototypes, or used directly for small lot production.This is an ongoing effort. The immediate following steps are to improve thelaminator design to achieve higher pressures and more flexibility, switch to working with commercially available CMC material systems, further investigate lasers that combine the advantages of the copper vapor laser and Diamond CO2 laser, and investigate low cost tooling materials which would provide more robust mandrels for the curved layer process than LOM paper. Also, the laminator size envelope will be expanded to near or equal to that of a standard LOM2030 machine (32" x 22").REFERENCES1. Klosterman, D., R. Chartoff, N. Osborne, G.

Also, the laminator size envelope will be expanded to near or equal to that of a standard LOM2030 machine (32" x 22").REFERENCES1. Klosterman, D., R. Chartoff, N. Osborne, G. Graves, "Automated Fabrication of Monolithic and Ceramic Matrix Composites via Laminated Object Manufacturing (LOM)," Solid Freeform Fabrication Proceedings, University of Texas at Austin, Austin, TX, Sept. 1997, pp.537-549.2. Klosterman, D., R.Chartoff, G. Graves, N. Osborne, A. Lightman, G. Han, A, Bezeredi, S. Rodrigues, "Automated Fabrication of Monolithic Ceramics and Ceramic Matrix Composites (CMCs) Using a Novel Rapid Prototyping Method," Proceedings of the 21 st Annual Conference on Composites, Advanced Ceramics, Materials, and Structures, The American Ceramic Society, Cocoa Beach, FL, Jan., 1998.3. Kalmanovich, G., '''Curved-Layer' Laminated Object Manufacturing," Solid Freeform Fabrication Proceedings, University of Texas at Austin, Austin, TX, Sept. 1996, pp.273-279.4. Lightman, A., G.

Kalmanovich, G., '''Curved-Layer' Laminated Object Manufacturing," Solid Freeform Fabrication Proceedings, University of Texas at Austin, Austin, TX, Sept. 1996, pp.273-279.4. Lightman, A., G. Han, "Laser Cutting of Ceramic Composite Layers," Solid Freeform Fabrication Proceedings, University of Texas at Austin, Austin, TX, Sept. 1996, pp.291-298.5. Klosterman, D., B. Priore, R. Chartoff, "Laminated Object Manufacturing of Polymer Matrix Composites," i h International Conference on Rapid Prototyping, University of Dayon and Stanford University, San Francisco, CA, March 31-Apri13, 1997, pp. 283-292.676ILLUSTRATIONSDesired Part ~Flat Layer LOM (9 layers)Curved Layer LOM (4 layers)= fully utilized material = wasted materialFigure 1: comparison of the flat layer and curved layer LOM processes for building a curved shell.

Due to the mode of building, the curved layer process will require fewer layers of material to build the same part.example of heater conforming to partvacuum pick-up and delivery I Deformable heater systemx-v laser control gantryRotatable feed tableprebuilt LOM-paper mandrelLOM pl8tformtFigure 2: curved layer LOM process schematic.677Silicone rubber frameFlexible heater filmFigure 3: curved layer bonding apparatus (prior to installation in LOM machine). The vacuum cups, which protrude though the structure normal to the heater surface, are not installed on the unit in this picture.Figure 4: curved layer, monolithic SiC body armor panel, immediately after LOM processing. The piece has been fully decubed and placed back on the LOM-paper mandrel for illustration.

The piece has been fully decubed and placed back on the LOM-paper mandrel for illustration. Notice the smooth surface and lack of stair steps.678SiC/SiC tape.~m:J.gCurved· Layer LOMDecubing !.__---'--J.-".-....,Reaction bonding 16000Co o o o••Full Pyrolysis 100°CPowder pressing: 100% cure, 50% pyrolysis 325°C, 60 psiFigure 5: overall process flow chart for fabricating net shape CMCs with curved layer LOM process.Figure 6: Re,lcti(m b()ndl~d SiC/SiC aircraft engine flame holders of various curvature, and one smail-sca.le monolithic SiC body armor panel.679Figure 7: Polished cross section of (left) eleven-layer SiC/SiC LOM composite prior to pyrolysis, and (right) after pyrolysis and reaction bonding.

From top to bottom the layers are: ceramic tape, 90° fiber layer, ceramic tape, 0° fiber layer, ceramic tape, etc.00 fiber layerFigure 8: Photomicrograph of reaction bondedSiC/SiC microstructure showing mating of various layers, infiltration efficiency, and fiber condition.680.

urved-Layer" Laminated Object Manufacturing®Gary Kalmanovich HelisysTM Inc.AbstractA process where a freeform solid object is built by bonding cut layers of material together is called a Laminated Object Manufacturing process. Until today, all of the LOM processes utilized straight forward planar cross-sections of the intended object for manufacturing. The "Curved Layer" project's objective is to perform the LOM manufacturing process on a non-planar cross section. Thus, the layers are going to be bonded together as non-planar surfaces providing additional strength to the built part as well as expedition of the built time for many objects. This the implementation of the project entails both software and hardware development manufacturing process.for1.IntroductionLaminated Object Manufacturing (LOMlM )is a process where a solid freeform object is manufactured with lamination of cut layers of sheet material.

Until today all of the (LOM) processes have been implemented using straight forward lamination of material cut flatly based on planar cross-sections. The "Curved-Layer" project's objective is to perform the LOM manufacturing process on a "curved," non-planar, cross-section. This is especially useful for thin, non-flat parts. This project entails numerous approaches and techniques that are not currently available commercially. Consequently, an innovative effort is required to perform a build of an LOM part on a non-planar surface. A substantial modification of the conventional (planar) LOM process is necessary in both software and hardware.2. Planar LOM ProcessThe planar process for laminated object manufacturing has been utilized by Helisys Inc. for rapid prototyping for some time. The most basic description of the process can be provided using a cutting and bonding cycle, see Figure 1.

for rapid prototyping for some time. The most basic description of the process can be provided using a cutting and bonding cycle, see Figure 1. It is important to outline the main stages of the current process in order to better understand what needs to be done for the "Curved-Layer" project.a) In the implementation of the current machine, the cutting is performed using a plotter system that transports a mirror and a focusing lens over the x-y envelope of a flat cross-section of the part corresponding to the [ z = current height ] plane. Everything, including the focusing lens, is fixed on the x-y plotting mechanism because the flat-layer is always parallel to the plotter's operational plane.b) Once a layer is cut, a new, virgin layer is bonded on top by a heater-roller. Note, the material has heat sensitive adhesive on the bottom side.

Note, the material has heat sensitive adhesive on the bottom side. The heater roller is flat in the y direction and is designed to apply equal pressure and temperature to the part as it rolls over it.®Laminated object manufacturing is a registered trademark of Helisys Inc. Torrance, CA This work was sponsored in part by a grant from DARPA.TM LOM and Helisys are trademarks used by Helisys Inc. Torrance, CA273)Laser CuttingNew Layer BondingFigure 1 Planar LOM processc) After bonding, the height of the part is measured and a new cross-section is computed by intersecting the part (.STL) file and the [ z = current height] plane.

After all the segments defined by the intersection of part facets with the height plane are identified, processing such as ordering, cross-hatching, beam-offset, among others are performed.d) When the necessary calculations are completed, the new material is cut, thus returning to the beginning of the cycle, i.e., return to a).is composed of components that in their design are inherently planar. The laser cutting, due to variable distance; the new layer bonding mechanism; the dependence on the part height measurement, because of reliance on part flatness; the planar cross-section, among all of the other 2D software components; and all the other LOM elements need to be rethought.Thecurrent planar LOM processFigure 2 Roller bonding3. "Curved-Layer" LOM ProcessThe "Curved-Layer," non-planar LOM process in its concept is almost identical to the regular, planar LOM process, i.e., compare Figure 3 to Figure 1.

"Curved-Layer" LOM ProcessThe "Curved-Layer," non-planar LOM process in its concept is almost identical to the regular, planar LOM process, i.e., compare Figure 3 to Figure 1. However, the two processes have to beLaser CuttingNew Layer BondingFigure 3 "Curved-Layer" LOM process274very different in their implementations. implications for both hardware and software are discussed.In this section,the relevant differences and their3.1 Non-Planar BaseThe planar LaM process uses a base constructed by adding flat layers of material to the metal platform. Afterwards, the part build cycle is begun on top of the base. The first major difference that the non-planar LaM process requires is the capability to form a base for the "Curved-Layer" build, see Figure 4. This is actually an intricate problem in itself Many circumstances such asFigure 4 "Curved-Layer" baseholes in the original object, high slopes, and etc.

This is actually an intricate problem in itself Many circumstances such asFigure 4 "Curved-Layer" baseholes in the original object, high slopes, and etc. may prevent the implementation of a non-planar part build because their presence in the base might interfere with the possibility of adding new layers of material. Therefore, such circumstances need to be eliminated in base creation itself Thus, the software for such base generation needs to have intelligence to identify and address the software developed uses some coarse heuristic approaches to such features. Currently, remove the undesired extrusions, etc. However, this method is not foolproo:f, though we have been able to create satisfactory bases for most object files we tried.Another new issue that exhibits itself is the base alignment issue.

However, this method is not foolproo:f, though we have been able to create satisfactory bases for most object files we tried.Another new issue that exhibits itself is the base alignment issue. Right now, we create the base by having the software construct a .STL file for this base and consequently building the base from the .STL file using the conventional, planar, LaM process. Because the base is created separately, prior to the non-planar build of the part, the base needs to be placed in perfect alignment with where the base perceives it to be. At this point, we do not think it to be practical to add new and very involved complexities of designing a closed loop system for automatic alignment.

At this point, we do not think it to be practical to add new and very involved complexities of designing a closed loop system for automatic alignment. However, a manual procedure needs to be worked out for performing this task.thecomputer'srepresentation for3.2 Non-Planar Cross-SectionIn order to identify the path required to cut on the recently bonded layer of material, we need to intersect the "Curved-Layer" with the object, i.e., identify the segments that constitute the intersection of the two, see Figure 5.Figure 5 A "Curved-Layer" intersecting an object (cube)275The mathematical "Curved-Layer" representation, in tum, needs to be a very accurate depiction ofthe actual top layer freeform. For planar layers the height measurement is taken after every new layer is laminated. Unfortunately, for non-planar layers there is a different height at every point.

For planar layers the height measurement is taken after every new layer is laminated. Unfortunately, for non-planar layers there is a different height at every point. Thus, either a closed loop system is needed to measure a layer's profile at every layer, where we will gauge the height at every point on the layer, or this can be accomplished with a very accurate open loop system, where the height at every point will be approximated with a mathematical model as the part grows. Just as the case with alignment, because a measurement system would bring in considerable extra complexity and become a major detraction from the actual problem at hand arid because if needed this feature it can be added later, we decided to proceed with an open loop approach at this time.3.2.1 Height grid representationWe looked at a number of candidates for the model layer. After appreciable consideration we decided to use a height grid system.

After appreciable consideration we decided to use a height grid system. In this system, we divide the xy-plane into an evenly spaced grid and store the height (z) values as elements of this grid, see Figure 6. This approach has the advantage of being much more elementary than most of the other models. Also, by using this model, the update and initial values for the non-planar layer are more fundamental and can be obtained using more tangible methods.to representthe non-planarFigure 6 Height grid representation3.2.2 Initial valuesThe initial values for the heights are first obtained by intersecting the triangles at the bottom of the object's .STL file with the vertical lines originating at the bottom of the grid, see Figure 6. Then, "foreign" values for the heights are filtered out using the same heuristic approaches that are used for the generation of the non-planar base generation.

Then, "foreign" values for the heights are filtered out using the same heuristic approaches that are used for the generation of the non-planar base generation. Note, this filter and the filter used for base generation are to have the same parameters in order to obtain an exact correspondence between the top layer of the base and the representation for the first layer.3.2.3 UpdateAfter a new successive layer has been added above the previous one, a new "Curved-Layer" representation needs to be obtained. This new layer is not a perfect replica of the old layer. Though, it Figure 7 Part growth should be a "smoothened" version of the previous layer, see Figure 7. Because we have no way to measure these new heights, an open loop method was developed to approximate the new, updated grid heights for this new layer.

Because we have no way to measure these new heights, an open loop method was developed to approximate the new, updated grid heights for this new layer. The approach chosen for this case consists of taking a point form the previous grid along with four of the original surrounding points on that grid; constructing four triangles originating from our desired point; lifting each of the triangles by the material thickness in the direction of their respective normal vectors; fitting a curved surface tangent in desired places to the four offset triangles (a third degree polynomial is used); and finally taking the z-height value on that polynomial at the desired location in the xy- plane, see Figure 8.276This approach for updating the layer representation due to an addition of an extra layer has advantages over other methods we the smoothing that this method can model considered because, results after successive addition of layers.

It also is superior to a 3D triangle based format of modeling curved layers, because this method does not grow in complexity. For this method, with each successive layer, the fineness of the grid remains exactly the same and thus the computational complexity is not increased, unlike most of the other methods we considered for which the complexity tends to grow by a factor with every layer. Moreover, in simple tests this method seems Figure 8 New layer model to perform very well in modeling the build growth even after a few hundred layers.3.2.4 A non-planar cross-sectional cutlayer and the object representation constitutes a non-planar cross-section, i.e., an object's periphery along which we would need to cut, see Figure 5. This intersection consists of segments that correspond to facet intersections, see Figure 9.An intersection between a non-planarAt times, the facet intersections lead to hard to handle, degenerate cases of triangle intersections, see Figure 9(c,d).

These occurrences are numerically very unstable and lead to problems with extra and missing segments. Though for planar cross-sections, these cases are extremely rare and they can be avoided in a vast majority of occurrences, Figure 9 Facet intersections unfortunately for non-planar case, this is much more likely to occur and not as easy to avoid. This has a potential to be a major problem for certain parts.3.2.5 Generalization of segment handling to 3DThe current LOM software for the planar process contains for handling and utilizing the collections of many routines segments in two dimensions, i.e., sorting, rearranging, beam offset, cross-hatching, etc. Most of these routines are just as much an absolute necessity for curved slicing as they are for planar slicing, e.g., arrangement of segments into continuous loops. Others, such as crosshatching are also very important.

Others, such as crosshatching are also very important. And finally, features such as the beam offset (to account for the thickness of the beam) are desirable and come with very little cost, assuming the other components are incorporated.We have extended the segment handling capability of the planarLOM software to 3D almost entirely, see Figure 10.Figure 10 Segment handling2773.3 Non-Planar Cutting (Focusing)As mentioned earlier, the distance from the focusing lens to the cutting surface varies for "curved layer" cutting. Because the plotter transporting the focusing lens is a 2D device and the surface is no longer flat, the cutting surface will be out of focus if the cutting is done in the conventional, planar manner.

This is another major difference between the two processes that need to be addressed.3.3.1 Platform compensationFigure 11 Non-planar cuttingOne approach to keep the laser constantly in focus is to move the platform supporting the part to compensate for the variability in height. This approach, in concept, does not require any hardware changes. Unfortunately, in practicality, this is not true because the currently used motion control _.... boards, as well as most motion control boards on the market, do not support circular arc motion in 3D. However, it is possible to implement the current motion using the current hardware if we stop the cutting motion every once in a while in order to adjust the height, see Figure 12. Nonetheless, this method is not the most practical because the motions of the platform are slow and are not always smooth. Moreover, the jitters resulting from platform motion can interfere with cutting.

Moreover, the jitters resulting from platform motion can interfere with cutting. Redesign of the platform system is costly, time consuming, and does not guarantee a satisfactory result.Ideal PathActual PathFigure 12 Platform compensation3.3.2 Large focal distance lensAnother method, to compensate for the variable part height, we are considering is to increase the focal distance of the lens, i.e., the focus region is extended by using a lens that focuses the beam further from itself, see Figure 13. The drawback of this approach is that the width of the beam at the focus spot is not as tight as it is with the original lens. As a result, it is not certain that we will be Figure 13 Large focal distance able to cut the material with this new lens.

As a result, it is not certain that we will be Figure 13 Large focal distance able to cut the material with this new lens. At this time, we are lens making the necessary hardware changes (to the lens housing, etc.) in order to try to cut with it at real time speeds.3.3.3 Controlled focal distancethe last method we are exploiting for this problem is the possibility of moving the focusing lens up and down so that the laser is kept in focus on the cutting surface. If the other methods are found not satisfactory then this method will be developed fully. The only drawback of this method is that the added weight of the focusing system will effect the balance of the plotter.

The only drawback of this method is that the added weight of the focusing system will effect the balance of the plotter. Thus, the plotter might have to be modified.Finally,Figure 14 Controlled focal distance278Of the three, this method is the most general with the fewest drawbacks, though it is the hardest to implement because it requires a mechanical, an electronic, and a software redesign effort.3.4 Non-Planar BondingThe bonding issue is of most concern to us and is the issue that seems to be the least open to workable solutions. It is not completely apparent how one is to deliver both heat and pressure uniformly to a loose sheet of material on a surface of a non-flat object so as to achieve smooth bonding with no ripples. Many approaches have been considered and rejected Figure 15 Bladder bonding due to one drawback or another.

Many approaches have been considered and rejected Figure 15 Bladder bonding due to one drawback or another. Weare still considering a few approaches, but are in the design stages for only one, though it too has many drawbacks.iL._...._...._..._..._..._..._..._..._..._...-4We are building a hot rubber bladder bonding mechanism. The bladder is to be filled and pressurized with circulating hot air. The rubber surface (siliconized rubber, so it can withstand high temperatures) comes in contact with the part by raising the part on the platform into the bladder. A previous non-pressurized prototype has produced some favorable results.3.5 Sheet FeedingBecause often it is much more accommodating to produce many materials in sheets as opposed to rolls, especially the materials we are considering for "Curved-Layer" building, the idea of a sheet feeder for the material is being considered.

(Thus far the LOM process has only been implemented with rolls of LOM material.) However, the mechanical design of such a mechanism would require an almost exact knowledge of many characteristics of the material in question. Thus far, the materials to be utilized are only being developed. Such characteristics as stiflhess, texture, lamination to the stack, and etc. are still not known. Moreover, if the material is too brittle, not stiff enough, or sticks to the stack too much, it might not be possible to implement a sheet feeding mechanism at all.4. SummaryThe "Curved-Layer" project to modify the LOM process from planar to non-planar build necessitates a substantial design effort. Many issues such as mechanical, electrical and software need to be addressed. In this paper we discussed our progress on many of these issues and what our concerns at this point in time are. Major components of the software effort have either been implemented or are in the implementation stages.

Major components of the software effort have either been implemented or are in the implementation stages. Though, some new software components will be the hardware components still necessary when certain hardware designs are finalized. Also, necessitate a substantial effort in their design and implementation. Weare going into uninvestigated territory and experimentation is needed before a final satisfactory product can be formed.279.

STOM FABRICATION OF HARD TISSUE RECONSTRUCTIVEFRAMEWORKSR. V. Nambiar, R. E. Jones and J. R. GomezThe University of Texas-Pan American, Edinburg TX 78539AbstractThe feasibility of fabricating custom frameworks for tissue regeneration utilizingthree-dimensional inkjet printing technology followed by slip casting was investigated. ACAD solid model of mold was created with structures within the mold cavity to providevariable, customized porosity in the cast framework. The polymer mold was printed usinga 3D ink-jet printer. The mold was then infiltrated by an aqueous suspension ofhydroxyapatite to produce porous frameworks that were tested for mechanical properties.IntroductionPorous materials have been used as non-resorbable intermediate structures forbone repair and when pore size is properly controlled such repairs can integrate into newbone, even permitting the development of vascularity [1].

A promising approach for hardtissue reconstruction is the integration of a custom mechanical framework with biologicalagents to promote development of the desired tissue [2]. Such a structure must havethree characteristics. It must fit the patient, thus requiring a unique structure for everyrepair. It must have specified mechanical properties to permit proper function duringregeneration. Finally, it must be able to carry the required precursor cells and/or growthfactors and direct their development. Thus it will require a controlled internal structurethat will probably not be uniform if it is to guide development of all of the subcomponents539of functioning bone or cartilage.

Thus it will require a controlled internal structurethat will probably not be uniform if it is to guide development of all of the subcomponents539of functioning bone or cartilage. These requirements make solid free form fabrication thebest hope for cost effective fabrication of the required structures.Research to date on rapid fabrication of tissue frameworks has utilized eitherselective laser sintering [3], laminated object manufacturing [4] or fused depositionmodeling [5]. In each case, the part must be sintered after the polymer binders are burnedout. The ink-jet based rapid prototyping technology proposed for this study is well suitedto the fabrication of precision frameworks for reconstruction. The system prints a layer ofbuild material using an ink-jet head and maintains accuracy in the z-direction by milling thelayer to a precise height before the next layer is deposited.

The system prints a layer ofbuild material using an ink-jet head and maintains accuracy in the z-direction by milling thelayer to a precise height before the next layer is deposited. The result is one of the mostaccurate and affordable commercially available solid free-form fabrication system. Modelswith pore feature size of 0.5 mm repeatability accurate to 0.05 mm have been produced.This process is not suitable for very large parts that may be required in orthopedicreconstruction, but it is capable of producing many of the common components needed formaxillofacial work.In this work, a CAD solid model of a mold is built with the inner surface createdusing imported surface data from a CT scan. Structures are created within the mold cavityto provide variable, customized porosity in the cast framework. The polymer mold isprinted using a 3D ink-jet printer. The mold is then infiltrated by an aqueous suspension ofhydroxyapatite mixed with binder.

The polymer mold isprinted using a 3D ink-jet printer. The mold is then infiltrated by an aqueous suspension ofhydroxyapatite mixed with binder. The part is dried then processed in an oven to burn outthe polymer binder and then sintered to produce porous framework.The ability of tissue growth into the framework is requires a well connected network of macropores.A number of different designs for the internal geometry were considered. A repeated hexagonalpattern of axial cylinders interconnected like a honeycomb was found to give the most uniform540mesh as shown in Figure 1. The pattern is repeated at uniform axial increments throughout thestructure. Cross section sizes ranging from 0.25 mm2 to 0.7 mm2 were used, the smaller sectionsshowing a higher rate of damage in processing. These structures in the wax mold will providepredictable sized pores in the cast structure.Figure 1.

These structures in the wax mold will providepredictable sized pores in the cast structure.Figure 1. CAD model of mold with internal structureProcessingHydroxyapatite powder (Aldrich) was mixed with water and compressed to form coupons.Samples of the coupons were analyzed in a differential thermal analyzer (DTA).The graph in Figure 2 shows the decrease in weight of the sample as thetemperature increases. The initial downward slope is attributed to the loss of water fromthe specimen up to 600C.541Hydroxylapatite - Weight % vs. Temperature10110099%98t h g i e W9796959402004006008001000120014001600Temp. (deg C)Figure 2. DTA weight loss analysis of HAThe graph in Figure 3 shows the temperature difference of the sample from thecontrol pan in the chamber due to exothermic or endothermic processes.

( d e g . C )Figure 4. DTA weight loss analysis of HA with binderH y d r o x y l a p a t i t e + B i n d e r - T e m p e r a t u r e D i f f e r e n c e v s . T e m p e r a t u r e2.521.5e c n e r e f f iD1.p m e T0.5002004006008001000120014000.5T e m p . ( d e g . C )Figure 5. DTA temperature analysis of HA with binderFigures 4 and 5 show the corresponding graphs for hydroxyapatite mixed with DuramaxB-1020 binder. Figure 4 shows the loss of water from the specimen at 500C. The graph in54316001600Figure 5 shows binder is burnt off at 700C. This establishes the burn off temperature at600C followed by sintering at 1100C, thus the temperatures for processing the specimenshave been identified.Molds with the outer body of the mold and the internal structures for the pores allin one piece were built using the Sanders ink-jet printing machine.

Hydroxylapatitepowder was combined with Duramax B-1020 organic binder and water to make a slurry.The slurry was poured into the mold and the water allowed to seep out of the base to formsolid ceramic specimens.The specimens fabricated above were all found to have cracks and internal cavities. Thesemake compression strength tests unreliable. The cause of these defects appears to beuneven drying. Experimentation to eliminate these defects by modifying the binder contentand controlling the humidity is continuing.SEM scans of the surface reveal as shown in Figure 6. that the sintering processleads to relatively smooth surfaces. This may indicate that the surface needs to be etchedto create micropores to enhance osteoinductivity.Figure 6. SEM picture of part surface 4000X magnification544ConclusionThe results of the studies conducted this year suggest that we can design molds forframeworks with specified pore size and distribution in a CAD system.

STOM, INTEGRATED, PNEUMATIC, ROTARY ACTUATOR FOR AN ACTIVE ANKLE-FOOT ORTHOSISRichard Remmers†, Doug Cook∗ , and Vito Gervasi∗†Bucyrus, Intl., 1100 Milwaukee Street, Milwaukee, WI, 53172, USA ∗ Rapid-Prototyping Research, Milwaukee School of Engineering,1025 N. Broadway Street, Milwaukee, WI, 53202, USAABSTRACTEnd-use objects produced via additive manufacturing (AM) are on the rise and newapplications in the fluid power industry are emerging. Recently, a custom, pneumatic, rotary actuator was been designed and additively manufactured for integration into an active ankle-foot orthosis that is being developed in the National Science Foundation’s Center for Compact and Efficient Fluid Power. All necessary plumbing, between the valves and vanes, is integrated into the additively-manufactured housing of the actuator; and, the silicone translating seals were vacuum-transfer molded using additively-manufactured molds and inserts.

This non- conventional actuator has more theoretical torque, and weighs less, than the off-the-shelf component that it replaced. Further development will reduce seal leakage, and optimize designs for additional mass reduction. Results-to-date are presented, in addition to several other examples of the growing use of AM in the fluid-power industry.1. INTRODUCTIONPurposeApplications of AM are expanding each year in a multitude of industries due to increasedAM awareness (through education), process and material improvements, and system cost reductions. The fluid power (FP) industry has used AM for NPD for well over a decade, creating prototype systems for form, fit, function and production. Now, end-use AM components and systems are becoming a reality and this research project demonstrates mass reductions and flexible reconfiguration options enabled by AM.

Now, end-use AM components and systems are becoming a reality and this research project demonstrates mass reductions and flexible reconfiguration options enabled by AM. The challenges faced on this project were to produce a reduced-mass rotary actuator for an Active-Ankle-Foot-Orthosis (AAFO) within a tight timeline with increased torque compared to the state-of-the-art off-the-shelf actuators available. This project also served to fulfill a masters capstone design thesis for the completion of MSE-ME. This paper has been prepared to share this example of AM technologies being leveraged directly and indirectly to solve a real-world problem in the FP industry.ScopeThe approach to solving this problem was initially to create a model using National Instruments LabView software to aid in establishment of a “working window” for size and performance of the actuator.

Next, an actuator was designed based on existing off-the-shelf actuators with the added degrees of design freedom offered by AM fabrication. And finally, the design was refined using several FEA packages to further reduce mass and increase housing stiffness. A combination of AM-based fabrication and conventional machining was used for this time-critical project. Other possible solutions were considered to generate the needed torque including mechanically-linked cylinders, motors, and a yet-to-be-disclosed fluid power actuator.Within the time constraints the “lower-risk” AM-based rotary actuator was chosen. The most critical considerations in choosing the path for this actuator included time, mass, torque and size. The timeframe was short at less tha 25 weeks and the mass needed to be less than one kilogram for the entire AAFO. The torque target was 6-10 N-m and the size needed to be reasonable, with a profile and compactness improvement over the current design.Figure 1.

The torque target was 6-10 N-m and the size needed to be reasonable, with a profile and compactness improvement over the current design.Figure 1. Prototype 1 Active-Ankle- Foot-Orthosis (AAFO), UIUC.BackgroundThe Center for Compact and Efficient Fluid Power (CCEFP) is a National ScienceFoundation (NSF) Engineering Research Center (ERC) based at the University of Minnesota. The CCEFP is a collaborative network of researchers, professors and industry partners working to advance the field of fluid power technology. One of the Center’s design platforms, designated Test Bed 6 (TB6), embodies research and design at the 10-to- 100-Watt-system range and is currently the lowest-powered test bed of the Center. TB6 currently features the study of Active Ankle Foot Orthosis (AAFO) technology. These are braces that help people who have lost mobility in their ankle(s) by providing force to assist toe-up and toe-down motion.

These are braces that help people who have lost mobility in their ankle(s) by providing force to assist toe-up and toe-down motion. Under the support of the Center for Compact and Effiecnet Fluid Power (CCEFP) the UIUC had constructed an AAFO using an off-the-shelf rotary actuator and the unit, shown in Figure 1, was functioning as intended with a few limiting factors including size, mass, and torque. The rotary actuator provided more rotation than needed, 90 degrees. The specifications for the replacement actuator were as follows:Mass less than 1 kg for entire AAFO • Rotation of greater than 25 degrees • Size and profile smaller than prototype I • Torque target of 6-10 Nm • Operating pressure of 0.35MPaEfforts on prototype II began during mid-summer of 2009. Clinical testing of the AAFOwas scheduled to begin in late 2009 and the timely replacement of the “bulkier” configuration was considered time-critical.

Clinical testing of the AAFOwas scheduled to begin in late 2009 and the timely replacement of the “bulkier” configuration was considered time-critical. To achieve the goal of upgrading the prototype I AAFO to meet the specifications conventional actuator fabrication techniques were not able to provide the addition of a “third” blade in the small actuator package needed to achieve the torque at the lower2a2b2c Figure 2. Fluid Power Actuators Grown Directly using Additive Manufacturing (2a- Wisner-2004, 2b-Zientarski-2009, 2c-Kang-2004).pressure and reduced mass. To address these design constraints AM was identified as the key enabling technology to deliver the actuator in the desired time-frame to meet specifications.The theory and knowledge-base informing the FP side of the project relied on basic fluidmechanics theory.

The AM side of this project drew form a general understanding of the capabilities and limitations of AM obtained from experts and literature as well as suppliers and material property data-sheets. A number of projects leveraging AM for fluid power components and systems have been performed at the MSOE Rapid Prototyping Center including several by former REU students. The first related project was titled, “Direct Manufacturing of Fluid Flow Devices in a Single Build” by Wisner (shown in Figure 2a) and a more recent project titled, “Additive Manufacturing of Fully Functional Fluid Power Components,” by Zientarski (Figure 2b). There are many examples of AM being used to create desirable FP actuator movement through the use of additive technologies including the micro-fluidic actuator shown in Figure 2c.

There are many examples of AM being used to create desirable FP actuator movement through the use of additive technologies including the micro-fluidic actuator shown in Figure 2c. This project leveraged the capability of AM to grow complex internal structure combined with the tight-tolerance capability of traditional material removal methods, specifically CNC milling and turning.Actuator DesignThe design goal was to reduce the number of individual components, and the size of the actuator, in order to reduce both mass and potential leak sites. The pneumatic diagram shown in Figure 3 illustrates the sub-system functions to meet the specifications for the AAFO. The grey shaded region includes those components integrated into the custom fluid-power system. The power source is regulated externally to provide two separate pressure levels, allowing for the generation of different torque outputs in either direction of rotation.

The power source is regulated externally to provide two separate pressure levels, allowing for the generation of different torque outputs in either direction of rotation. Actuation is controlled by two, solenoid, directional-control valves; and, the pressure in the actuator is released to atmosphere through an adjustable flow-control valve and silencer. This setup provides more control over rate-of-motion and exhaust noise.50 psi30 psi3: Pnuematic Diagram of Prototype System Figure 3: Pnuematic Diagram of Prototype Systemutstanding features of this actuator is the use of blind, internal passages One of the most outstanding features of this actuator is the use of blind, internal passages utstanding features of this actuator is the use of blind, internal passages are quite common and l volumes. Actuators with two rotor blades are quite common andto connect different control volumes.

Actuators with two rotor blades are quite common andto connect different control volumes. Actuators with two rotor blades typical methods of connecting the chambers use conventional machining operations, e.g. a hole typical methods of connecting the chambers use conventional machining operations, e.g. a hole typical methods of connecting the chambers use conventional machining operations, e.g. a hole complicated the typical two- center shaft. Adding a third blade complicated the typical two drilled tangentially through the center shaft. Adding a third bladeFigure 4: Actuator Housing Channels Figureblade form, necessitating an alternate to be connected and supplied the same pressure, the red sections needed n alternate approach. As shown in Figure 4, the red sections neededlikewise for the blue sections.

As shown in Figure 4, the red sections neededlikewise for the blue sections. lied the same pressure -- likewise for the blue sections.One design path would have been to attempt drilling three holes in the center shaft that One design path would have been to attempt drilling three holes in the center shaft that One design path would have been to attempt drilling three holes in the center shaft thatat a parallel location along the shaft. This met in the center at one location and three more holes at a parallel location along the shaft. This met in the center at one location and three more holes 4 supplied by the manifold and the d would have the top two chambers pictured in Figure 4 supplied by the manifold and the method would have the top two chambers pictured in other two chambers would be connected through the top chamber.

other two chambers would be connected through the top chamber.The preferred design utilizes the unique capabilities of AM to create channels that could The preferred design utilizes the unique capabilities of AM to create channels that could The preferred design utilizes the unique capabilities of AM to create channels that couldonly be generated via an additive process. There is no line circumferential channels that connect all of the chambers to their respective valves, and allow for circumferential channels that connect all of the chambers to their respective valves, and allow for circumferential channels that connect all of the chambers to their respective valves, and allow for more uniform pressurization. Figure volumes (blue channels are hidden behind red). volumes (blue channels are hidden behind red).sight for a tool to create these ted via an additive process.

Figure volumes (blue channels are hidden behind red). volumes (blue channels are hidden behind red).sight for a tool to create these ted via an additive process. There is no line-of-sight for a tool to create thesetch the control Figure 4 shows these channels colored to match the control5 are the design revisions of the actuator. Clearly visible is the reduction Shown in Figure 5 are the design revisions of the actuator. Clearly visible is the reduction 5 are the design revisions of the actuator. Clearly visible is the reduction in mass and component integration of the design. Revision 4 was the final design and component integration of the design. Revision 4 was the final design. Threaded and component integration of the design.

Revision 4 was the final design and component integration of the design. Revision 4 was the final design. Threaded and component integration of the design. Revision 4 was the final design brass inserts were placed with epoxy into the housing in several areas where fasteners required brass inserts were placed with epoxy into the housing in several areas where fasteners required brass inserts were placed with epoxy into the housing in several areas where fasteners required significant torque.Custom Actuator Design Revisions and Final Design (Revision 4). Figure 5. Custom Actuator Design Revisions and Final Design (Revision 4).Actuator ConstructionAs mainstream as stereolithographyhas become, there are limitations to the as stereolithography (SLA) has become, there are limitations to thel tolerances that can be achieved. Some of the features of the actuator required tight l tolerances that can be achieved.

Some of the features of the actuator required tight l tolerances that can be achieved. Some of the features of the actuator required tight dimensional tolerances that can be achieved. Some of the features of the actuator required tight tolerances. For this reason, these areas had material (machine stock) added to them so that they tolerances. For this reason, these areas had material (machine stock) added to them so that they tolerances. For this reason, these areas had material (machine stock) added to them so that they could be machined to tolerance after the part was olored red in the Figure 6 could be machined to tolerance after the part was grown.

The areas colored red in the were machined to improve the surface finish and tolerance of those surfaces as they are critical to were machined to improve the surface finish and tolerance of those surfaces as they are critical to were machined to improve the surface finish and tolerance of those surfaces as they are critical to operation of the actuator.6: Machined Areas of AM Housing Component Figure 6: Machined Areas of AM Housing ComponentThe seal located between the actuator housing and the babetween the actuator housing and the back plate was created using a s created using a 7, from a sheet of Buna-Nlaser cutting system.

The laser cut out a 70A durometer rubber.cut out a the seal, shown in Figure 7, from a sheet of BunaFigure 7: Seal7: Seal Between Back Plate to Actuator HousingThis method of making seals was found to work well for static seals where the faces cut for static seals where the faces cut This method of making seals was found to work well top and bottom faces of by the laser are not critical for sealing. Using only the as-manufactured top and bottom faces of by the laser are not critical for sealing. a rubber gasket sheet provided anadequate seal between the back plate and the actuator housing.

Using only the as-manufactured top and bottom faces of by the laser are not critical for sealing. a rubber gasket sheet provided anadequate seal between the back plate and the actuator housing. n adequate seal between the back plate and the actuator housing.nging part of this design and construction was the fabrication of the The most challenging part of this design and construction was the fabrication of the nging part of this design and construction was the fabrication of the 8) and in the actuator dynamic seals that reside in grooves on the three rotor blades (Figure 8) and in the actuator dynamic seals that reside in groove housing resulting in a moving seal on the rotor shaft. Their function is to form a seal betw housing resulting in a moving seal on the rotor shaft. Their function is to form a seal between housing resulting in a moving seal on the rotor shaft.

Their function is to form a seal betw housing resulting in a moving seal on the rotor shaft. Their function is to form a seal between housing resulting in a moving seal on the rotor shaft. Their function is to form a seal betw the rotor blade (or shaft) and the housing surfaces to maintain a pressure differential as needed to the rotor blade (or shaft) and the housing surfaces to maintain a pressure differential as needed to the rotor blade (or shaft) and the housing surfaces to maintain a pressure differential as needed to generate torque output while at the same time providing a sliding seal within the housing. The generate torque output while at the same time providing a sliding seal within the housing. The generate torque output while at the same time providing a sliding seal within the housing.

The generate torque output while at the same time providing a sliding seal within the housing. The generate torque output while at the same time providing a sliding seal within the housing. The challenge, then, is that the seals must: be rigid enoug h to resist deformation at maximum challenge, then, is that the seals must: be rigid enough to resist deformation at maximum operating pressure; slide with minimal friction across the sealing surfaces; and, conform to any operating pressure; slide with minimal friction across the sealing surfaces; and, conform to any operating pressure; slide with minimal friction across the sealing surfaces; and, conform to any imperfections.Figure 8: Three-bladed Rotor and Shaft. FigureThe initial attempts to create the dynamic seals were made using the same laser-cuttingtechnique used for the housing-to-back-plate seal. It was desired to use a low-durometer rubber for these seals so they would be more likely to conform to the surfaces.

It was desired to use a low-durometer rubber for these seals so they would be more likely to conform to the surfaces. The laser cutter could cut shore 70A rubber; but it would simply melt rubber with shore hardness less than 50A. No usable seals could be made out of the lower durometer rubber using the available laser cutter system. The shore 70A rubber was used for many attempts to make dynamic seals. Due to the width of the laser beam, there is a degree of inaccuracy in converting the CAD file to a seal; this is also affected by the precision of the linear stages and focus of the laser beam. The laser cut creates an irregular finish on the cut faces that resulted in leakage.Several sets of seals were made with varying widths. The goal was to find the proper thickness of seal that would create the optimal tradeoff between sealing and friction. On one occasion, when testing these seals the actuator did move, albeit with very little torque.

On one occasion, when testing these seals the actuator did move, albeit with very little torque. This was an encouraging moment; however, on the following day, the lubricant used in the actuator had been absorbed by the rubber causing it to swell. The actuator essentially locked up.A superior method of forming these insert seals was later used, again leveraging additivemanufacturing. A single-cavity silicone-transfer-mold was constructed from tool steel, and included a receptacle for an AM mold insert that contained the geometry of the seals, as well as the mold gating. Figure 9 shows the mold design, with the top portion rendered transparent. Tool steel was chosen because it was readily available in plates, and was already ground and lapped to an appropriate surface finish for the mold. Steel is also durable enough to withstand the frequent tightening and loosening of the bolts that hold the mold halves together.

Steel is also durable enough to withstand the frequent tightening and loosening of the bolts that hold the mold halves together. The AM mold insert (shown to the lower right in Figure 9) was thin, at 1.6 mm (1/16 in), to allow rapid build-times, therefore, multiple iterations could be made if necessary to adjust size and design to achieve a functional seal. A very low (shore10A) durometer silicone was cast to produce a set of low durometer seals.Figure 9: Steel Transfer Mold and AM Mold InsertOver-Molded SealsAdditionally, considerations were made for over-molding the blades of the actuator rotor with an elastomer. The lower shear loading at the blade tip, relative to the root, afforded a radial,tapered reduction in blade cross-sectional area, thereby increasing the area available for an elastomeric seal.

The lower shear loading at the blade tip, relative to the root, afforded a radial,tapered reduction in blade cross-sectional area, thereby increasing the area available for an elastomeric seal. Once again employing additive manufacturing, a silicone-transfer mold was fabricated using the Watershed™ resin, as was a prototype of the proposed rotor geometry (Figure 10). The transparency of this material, achieved with moderate sanding and buffing, allowed for inspection of the silicone flow; so that, trapped air, and any other potential defects, could be addressed before the material cured.Figure 10: Vacuum-transfer Mold for Overmolding the Blades of the Actuator Rotor with an Elastomeric-seal Geometry.Results and DiscussionThrough mass- and size-reducing design, the custom actuator is significantly smaller,lighter and more compact than the original prototype.

It is difficult to give a direct comparison of the first prototype actuator to the new prototype actuator because the new actuator incorporates many components in its assembly.

However, Table 1 shows the contrast in mass, along with the performance specifications of the prototype I actuator alone to the custom actuator.Table 1: Revision ComparisonPrototype ICustom ActuatorComponentsMassSMC Rotary Vane Actuator CRB2BW40- 90D-DIM00653 (alone, no valves/lines/etc.) 390 gRotary Vane Actuator, two directional control valves, two flow control valves, orthotic strut brace, electronics board 372 gTorque5 N-m at 0.35 MPa4 N-m at 0.35 MPaRotation Angle90°55°Figure 11a and 11b illustrate the SMC actuator and custom actuator (with the additionalcomponents listed in Table 1 installed, less the electronics board).(a) (b)Figure 11: Revision One and Revision Two ActuatorsGraph 1 illustrates the performance of the SMC actuator compared to the new actuator.Notable here is the drop-off in torque as pressure increases to 0.35 MPa. the drop-off in torque is attributed to friction and internal leakage of the dynamic seals.

the drop-off in torque is attributed to friction and internal leakage of the dynamic seals. Improvements in sealing were achieved during this stage of the project but an effort to optimize the geometry and durometer of the dynamic seals to achieve the theoretical 7 N-m was not carried out. Despite these challenges, the actuator has successfully demonstrated the use of AM to create a functional fluid power component. It has been tested up to 3,000 cycles. After more refinement of the seals, this actuator will be tested more extensively and the dynamics of the system will be identified to be used in developing the control system for the AAFO. Shown in Figure 12 is the orthosis with the actuator integrated into the unit.Actuator Revision ComparisonsActuator Revision Comparisons76m] N- [ e u q r o T5432MSOE Rotary1SMC CRB2BW40000.050.10.150.20.250.30.350.40.45Pressure [MPa]Graph 1: Revision One and Revision Two ActuatorsFigure 12.

Assembled AAFO with Figure integrated Actuator. integrated Actuator.shows the result of the first over-molded seal trial, using Polydimethyl While the fit of the assembly was good, changes in the performance requirements of (PDMS). While the fit of the assembly was good, changes in the performance requirements of While the fit of the assembly was good, changes in the performance requirements of this actuator have prevented us from testing the functionality of the over molded seals. The next this actuator have prevented us from testing the functionality of the over-molded seals.

The next this actuator have prevented us from testing the functionality of the over-molded seals. The next generation must provide more torque, and, therefore, must handle higher pressures to remain generation must provide more torque, and, therefore, must handle higher pressures to remain generation must provide more torque, and, therefore, must handle higher pressures to remain compact.Figure 13 shows the result of the firstPolydimethyl-siloxaneFigure 13: Over-molded silicone seals on the (center). The tapered blade withinone seals on the actuator rotor (left) and the fit into the housing actuator rotor (left) and the fit into the housing , being flexed (right).

The tapered blade withinone seals on the actuator rotor (left) and the fit into the housing actuator rotor (left) and the fit into the housing , being flexed (right). within can be seen in contrast to the silicone, being flexed (right)he torque requirements of this compact actuator necessitate the use of high-strength he torque requirements of this compact actuator necessitate the use of high The torque requirements of this compact actuator necessitate the use of highrotor’s shaft and blade roots. Future prototypes may integrate the two materials, e.g. steel, for the rotor’s shaft and blade roots. Future prototypes may integrate the two rotor’s shaft and blade roots. Future prototypes may integrate the two aterial (metal and ltimately, as additive technologies progress, such multi-material (metal and materials. Ultimately, as additive technologies progress, such multi rotors might be fabricated in one process.

Ultimately, as additive technologies progress, such multi rotors might be fabricated in one process. elastomer) rotors might be fabricated in one process.ConclusionsIt has been demonstrated that additive manufacturing is useful for creating functional It has been demonstrated that additive manufacturing is useful for creating functional It has been demonstrated that additive manufacturing is useful for creating functional fluid power components and tooling for creating seals for said components. AM has made a fluid power components and tooling for creating seals for said components. AM has made a fluid power components and tooling for creating seals for said components. AM has made a tremendous impact in changing the methods used to produce many custom personal products tremendous impact in changing the methods used to produce many custom personal products.For this application, plumbing and multiple components were integrated into a single housing.

AM potential has yet to be fully utilized in many engineered components. Through the use of AM, the disparity between an ideal design and a feasibly manufactured component can be reduced.Future WorkConsideration of higher pressure actuators has begun and future actuator operatingpressures may increase to greater than 500psi and air may be replaced with hydraulic fluid.AcknowledgementsThe authors would like to thank the National Science Foundation (NSF) and the Center for Compact and Efficient Fluid Power (CCEFP) for funding and support of this Research. This material is based upon work supported by the National Science Foundation under Grant No. EEC-0540834. Alex Shorter is recognized for his work on the first AAFO prototype that created the opportunity for this project. Also, Keith Nordgulen at SMC Milwaukee is acknowledged for his help.

EEC-0540834. Alex Shorter is recognized for his work on the first AAFO prototype that created the opportunity for this project. Also, Keith Nordgulen at SMC Milwaukee is acknowledged for his help. Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the views of the National Science FoundationReferences3D Systems “Viper™ SLA® system” [Internet, WWW, PDF] Available: Available in .PDF format; Address: http://www.3dsystems.com/products/sla/viper/ViperSLA_%20US_Engl_rev0806.pdf [Accessed: 25 May 2010].CCEFP (2009) “Transforming Fluid Power through Research | Center for Compact and Efficient Fluid Power.” [Internet, WWW]. Available: University of Minnesota Website; Address: http://www.ccefp.org/research. [Accessed: 23 July 2009].DMS Somos. (2007). “Watershed XC 11122” [Internet, WWW, PDF] Available: Available in .PDF format; Address: http://www.interpro-rtc.com/Somos11122_XC.pdf.

We tailor this procedure to a patient who will undergo a mandibulectomy due to cancer infiltration of the jaw. The patient CT scan data was used to generate a 3D representation of the patient’s skull, before the corrupted mandible was extracted. We examined two approaches based on classical symmetry matching and digital reconstruction of the defect to form the final model for printing. The final designs were then 3D printed and assessed for efficacy against a patient specific representative model of the skull and maxilla, where the final optimised design was found to provide an excellent fit. Ultimately, this technique provides a framework for the design and optimisation of a patient specific whole mandible implant.1 IntroductionThe minimisation of time for surgical procedures and patient recovery are highly pivotal in reducing financial burdens on healthcare providers and patients, whilst also improving patient outcomes.

This goal has led to many recent advancements and innovations in medicine, particularly in the area of patient specific treatment options. Disruptive technologies, such as additive manufacturing and digital 3D modelling have positively impacted areas of pre-operative planning and treatment [1, 2], and have led to the creation of a patient specific assistive and implantable devices, such as surgical resection guides [3-5], planning models [3, 6, 7], anatomical teaching aids [8, 9] and prosthetics [10-15]. More recently, the US Food and Drug Administration (FDA) have increased their approval of 3D printed implants under the 510k (pre-market notification) approval system, which permits the use of additive manufactured parts in routine and complex surgical procedures [16].

It is suspected that recent approvals and the wider acceptance of additive manufacturing is likely1708the result of increasing maturity of this technology, which has allowed the direct processing of an increasing number of biocompatible materials with ever increasing printing precision.A mandibulectomy is the removal of large sections of the lower jaw and is performed following either physical trauma or severe disease infiltration (cancer, etc) [17, 18]. Current surgical rehabilitation is performed using a guide plate, screw fixtures and grafted donor bone tissue (hip/femur) inserted into the void space to support the jaw. However, such methods are prone to short term failure (approx. ≤1 yr.), guide plates are predominantly manipulated during surgery to approximately match the patient’s contours and treatment arguably does not provide a solid base to allow osseointegration of a patient’s bone tissue.

It has been more recently found that fixation methodologies traditionally used are leading to complications such as mandibular fractures [6, 19, 20] resulting from fixation in structurally compromised bone tissue. Therefore there is increasing evidence to support the use of whole mandible replacement implants.Figure 01: A summary of the process chain for the construction of the complete mandible implantIn this study we assessed the potential of creating custom replacement mandible implant following a mandibulectomy. The implant was created using patient specific medical imaging data (CT DICOM) alongside state-of-the-art 3D design/manipulation software, before realising the implant using low-cost Fused Deposition Modelling (FDM) to assess finish the model (Figure 1). It is noted that the FDM printed models are for preliminary verification purposes and would not be used as the final implantable device.

It is noted that the FDM printed models are for preliminary verification purposes and would not be used as the final implantable device. This case study was performed on a patient who presented with cancer infiltration through the entire right-side of the mandible. CT scan images were used to generate a three-dimensional representative model of the patient’s corrupted mandible, before an initial full jaw replacement implant was constructed. Mandible replacement devices have been attempted previously, however this work differs from previously reported designs as we attempt to retain the major contours of the patient such that1709the implant retains the original facial structural contours of the patient. This contrasts favourably to other such examples of this form of treatment where the replacement mandible is generally much smaller that the original bone structure or require bone grafts for reconstruction.

We examined and compared the efficacy of reconstruction using anatomy mirroring and by direct mandible digital reconstruction. Ultimately, this work provides a framework for the design criteria to create an optimised patient specific, full mandible implant. The techniques used in this work could equally be applied to alternative patient anatomies, providing a generic methodology for implant design. Such work will provide a powerful tool for clinician’s to overcome the shortcomings of typical fixation plate based treatment option, thereby improving overall patient outcomes.2 Methodology2.1 Anatomical Data ModellingRepresentative models of the patient’s anatomical data were constructed based on Digital Imaging and Communications in Medicine (DICOM) data from CT scans using a Somotom scanner (Siemens Healthcare, Germany). In the DICOM format, the data is presented as a series of slices through the patient’s anatomy, each approximately 0.6mm thick.

In the DICOM format, the data is presented as a series of slices through the patient’s anatomy, each approximately 0.6mm thick. The software package Mimics (Materialise, Belgium) was used to compile the DIOCM data into axial, sagittal and coronal planes. Using Mimics, the various greyscales can be selectively marked to isolate a particular anatomical tissue type (i.e. bone, muscle, etc) and from this to construct a 3D model of this tissue type. The Mimics software has a series on inbuilt thresholds which allow for the rapid isolation of anatomical tissues and the inbuilt thresholds for bone were used to construct the final model, in additional to manual region growing.2.2 Implant DesignThe model created in Mimics was exported to 3-Matic (Materialise, Belgium) for further processing and construction of the mandible model. Prior to any operations, the CT scan extracted model underwent an initial phase of error correction.

Prior to any operations, the CT scan extracted model underwent an initial phase of error correction. This is required to remove excess data point, duplicate triangles, inverted normal, filter small sub shells and unify larger shells into a single model, allowing the model to be fit for 3D printing. The software was then used to segment the skull model to isolate the jaw. In this study two distinct approached were examined whereby the mandible was reconstructed using a mirroring of the patients healthy portion of the mandible and also by direct reconstruction of the cancerous section. Both approaches could potentially offer a high quality matching of the patient’s anatomy, but it was unclear which technique would be superior and so both approaches were assessed.2.3 Implant Prototyping and ManufacturingIn this study, the various design iteration of the jaw model were realised using fused deposition model (FDM) printing in ABS plastic.

FDM printing is a low cost technique for part production, and evaluation prints were realised using a Zortrax M200 (Olsztyn, Poland) printer.1710This methodology allowed for the rapid printing of the proposed implant design iterations to evaluate model build integrity and patient fitment/sizing against the original jaw anatomy.3 Results3.1 Anatomical ModellingUsing the inbuilt threshold function within Mimics (Bone CT- 226-3071), the bone tissue was predominantly isolated. However several features of the soft tissue were also captured at this greyscale value, in additional to several uncaptured data points of the bone. Therefore additional manual processing of the model was required to remove unwanted data point and to add missing bone data point in the CT scan slices across all three dimensional planes.

Therefore additional manual processing of the model was required to remove unwanted data point and to add missing bone data point in the CT scan slices across all three dimensional planes. Once this had been achieved to a satisfactory point the highlighted data set was then converted into a representative model of the patient’s wider skull, which contained the corrupt mandible.Figure 2: An image of the various planes of the patients CT scan data and the constructed model of the isolated bone tissue segment, containing the corrupt mandible.This constructed model comprised of a segment of the overall skull, the maxilla and the mandible. Primarily, the mandible was the area of interest in the jaw reconstruction, however inclusion of the skull and maxilla and realisation into a 3D printed model allowed for analysis of the jaw size and fit compared to the original anatomy of the patient.

Figure 2 shows the patient’s skeletal data and the resulting model that was formed for the mandible. Despite the automated threshold function and manual data point addition/removal, some elements can be missing, in additional to stray noise being present in the final digital model. In such instances a wrapping function can be performed improve the final quality of the model and without adjusting the size threshold of the model. Following wrapping the model was visually inspected1711to check that the major features were adequately rendered, before the model is exported as an STL file. The model is then loaded into 3-Matic to perform error corrections, segmentation, reconstruction of the corrupt region of the mandible and additional post processing.3.2 Implant DesignWhen the initial model is imported into 3-Matic error corrections are initially performed to ensure the integrity of the model prior to implant production.

Next the mandible was isolated to assess the degree of cancer infiltration in the patient. Figure 3a) shows a transparent digital model of the mandible where the cancerous region can be seen as a large cavity. At the extremities the cancerous region was estimated to be 25mm in the vertical and 50mm in the lateral direction.Two distinct approaches were examined to ascertain the best methodology for the implant production, and which comprised reconstruction through mirroring of the uncompromised area and direct reconstruction of the cancerous jaw segment. To assess the fit of the resulting mandible design, a model of the patient’s maxilla and temporomandibular joint was created from the patient’s DICOM data and printed using an FDM printer.

To assess the fit of the resulting mandible design, a model of the patient’s maxilla and temporomandibular joint was created from the patient’s DICOM data and printed using an FDM printer. Additionally, the corrupt mandible model was printed to compare against the corrected mandible model.Figure 3: a) The original corrupt mandible, b) outcomes of the symmetry based approach for mandible reconstruction and c) illustration of the anatomy mismatch of the final model against the original mandible.17123.2.1 Anatomy MirroringThe main area of cancerous infiltration was restricted to the left portion of the patient’s mandible. Therefore the mandible was divided into two sections and the right half mirrored to construct the final implant. When attempting to section the mandible, the teeth were retained to help guide the location for sectioning.

When attempting to section the mandible, the teeth were retained to help guide the location for sectioning. Upon closer examination of the mandible there were several natural contours in the front of the jaw which could act as a guide point for the sectioning process. Several locations were examined as the reference point to produce the complete implant, with promising results, as shown in figure 3b). Cross referencing of the original compromised section of the jaw revealed that due to asymmetry present in the original jaw, the location of the ramus, condylar process and coronoid process were all out of alignment (Figure 3c). Such misalignment would likely result in placement issues during implantation as well as compromising post-surgical rehabilitation and movement. Several attempts to re-adjust the spatial orientation of the mirrored mandible segments were unsuccessful in achieving a matching orientation to the original mandible anatomy.

Several attempts to re-adjust the spatial orientation of the mirrored mandible segments were unsuccessful in achieving a matching orientation to the original mandible anatomy. Based on these results, it was determined that anatomy mirroring in these particular circumstances of execution was not a robust technique for digital reconstruction of the corrupted mandible.Figure 4: a) Stages of the segment reconstruction strategy and b) The final model of the mandible.3.2.2 Segment ReconstructionIn the segment reconstruction approach, the corrupt portion of the mandible was isolated and reconstructed manually to create a solid portion that matched the original contours of the patients jaw. The advantage of this approach is that the majority of the shape and1713geometry of the original mandible are preserved. Initially, the mandible model was made transparent to assess and visualise the extent of the cancer infiltration, as shown in Figure 3a).

Initially, the mandible model was made transparent to assess and visualise the extent of the cancer infiltration, as shown in Figure 3a). This methodology allowed for the complete visualisation of the cancerous growth, which aided the digital reconstruction process. This compromised region of the mandible was segmented for additional processing.To aid the design process, the mandible was segmented into two halves about an approximate midpoint, to anatomy mirroring approach. The uncompromised half of the mandible was then mirrored and used as a reference template for the reconstruction of the segment, to achieve the desired thicknesses and surface contours. The segmented model in its native form comprised a hollow part with openings at both sides of the lateral sections of the mandible.

The segmented model in its native form comprised a hollow part with openings at both sides of the lateral sections of the mandible. It was also noted that the cancerous growth had caused for an expansion of the surrounding jaw line and so in addition to removal of the hollow sections, the segment required reworking to this lateral displacement. Several trimming and smoothing operations were performed to bring the side of the segment back to a similar dimension to that of the mirrored reference. It was additionally noticed that surface contours of the uncompromised region contained a variety of holes, beyond the natural cavities containing the mental nerve, which were required to be filled to ensure the maximum integrity of the final model. These holes were present on the patient’s original CT scan data and so reduced the holes/cavities being the result of digital artefacts during the modelling process.

These holes were present on the patient’s original CT scan data and so reduced the holes/cavities being the result of digital artefacts during the modelling process. This potentially implied that the mandible was structurally compromised and so should a traditional fixation plate approach be employed this could result in mandibular fracturing. These findings further confirmed the necessity for a total mandibular replacement. The final segment and its placement in the sectioned mandible is shown in Figure 4, where it can be seen that overall a very good reproduction of the original symmetry could be achieved.

The final segment and its placement in the sectioned mandible is shown in Figure 4, where it can be seen that overall a very good reproduction of the original symmetry could be achieved. This final model also matched the original locations of the ramus, condylar process and coronoid process due to only working on the cancer infiltrated region of the mandible, and therefore this reconstruction technique was considered the superior approach for this specific case study.in a similar manner3.3 Mandible 3D PrintingWhen constructing models based on data ascertained from medical imaging data, there could be facets of the model, such as low thickness tolerances in regions of the part, which could lead to compromised structural integrity. Given the complexity of such models, it may not always be possible to perform adequate thickness analysis and Finite Element Analysis (FEA) to validate mechanical properties.

Given the complexity of such models, it may not always be possible to perform adequate thickness analysis and Finite Element Analysis (FEA) to validate mechanical properties. Additionally, as such parts are intended as an implantable device, a rigorous and robust evaluation procedure is desirable. Therefore a preliminary low-cost polymer printing approach can help provide qualitative validation of a parts structural integrity, before the final, more costly printing in metallic materials.The final model was then 3D printed initially on an FDM printer to assess the structural integrity of the final models and also to ensure correct dimensional accuracy of the jaw against both the corrupt mandible and partial skull/maxilla models. It was found that the model printed very well with no perceivable issues with its structural integrity.

It was found that the model printed very well with no perceivable issues with its structural integrity. Figure 5 shows the printed corrupt and reconstructed mandible, alongside the reconstructed mandible implant placed into a representative model of the skull/maxilla. It can be seen that the reconstruction approach of the jaw was highly effective at reproducing the patients original jaw line, with minimal impact1714to the overall structural contours. The resulting mandible models proved an equally excellent fit into the representative model of the patient skull, as can be seen in Figure 5. These qualitative tests confirmed the efficacy of the design and the structural integrity when realising the mandible model by 3D printing processes.

These qualitative tests confirmed the efficacy of the design and the structural integrity when realising the mandible model by 3D printing processes. We are therefore confident that the design could be equally rendered by SLM processed in titanium, a certified material for medical implantation, and this will be the subject of future studies.Figure 5: Representative FDM models of a) the original cancerous mandible, b) the reconstructed mandible implant and c) the developed implant placed onto a model of the patient’s skull and maxilla.4 ConclusionsIn this study we have examined the use of medical CT scan data to reconstruct a digital model of a patient’s skeletal anatomy, and from this, a full replacement mandible to treat a cancerous infiltration. The final model and the patients surrounding skeletal anatomy was then realised using a FDM 3D printing process, to evaluate the size and geometric precision of the final implant.

The final model and the patients surrounding skeletal anatomy was then realised using a FDM 3D printing process, to evaluate the size and geometric precision of the final implant. Investigating various design techniques, it was found that a mandible reconstruction approach offered superior fit as compared to the anatomy mirroring approach. Ultimately, this study demonstrates an optimised approach for patient specific mandible reconstruction that could have significant potential for use as a strategy in patient specific, custom implant design.1715AcknowledgementsWe would like to thank the School of Engineering at Deakin University for providingfunds and resources for this pilot project.References1.2.3.4.5.6.7.8.9.10.11.12.13.14.15.Rengier, F., et al., 3D printing based on imaging data: review of medical applications. International Journal of Computer Assisted Radiology and Surgery, 2010. 5(4): p. 335-341.

STOMISED LAYER DEPOSITION FOR CHEMICAL REACTOR APPLICATIONSJ. Singh a, C. Hauser a,P.R. Chalkera, C. J. Sutcliffea,A.T. Clare a and M. Dunschen b. a Department of Engineering, University of Liverpool, UK b FreeSteel, Liverpool, UKAbstractThis paper discusses the development and application of an adaptive slicing algorithm for use with Digital Light Processing (DLP) for the manufacture of micro chemical reactors. Micro reactors have highly complex constructions and DLP has a proven ability to deliver features at the micro level with high accuracy. However, DLP fails to provide a truly smooth profiled surface finish which could influence fluid flow through entrance and exit apertures and along snaking micro channels. Ensuring smooth surfaces will minimise energy losses in the fluid flow path. Generally, layer based manufacturing techniques incur a trade off between build time and resolution.

Ensuring smooth surfaces will minimise energy losses in the fluid flow path. Generally, layer based manufacturing techniques incur a trade off between build time and resolution. The algorithms used in this study attempt to mitigate this to some degree by calculating locations where high resolution is required through surface profiling techniques and adjusts the layer thickness accordingly. It is proposed that this adaptive layering technique may improve surface roughness and reduce friction related energy losses along micro channels within chemical reactor applications.IntroductionMicro reactors are devices used for heat and mass transfer. This is achieved by bringing two fluids in direct or indirect contact in confined “micro channels” that essentially have lateral dimensions less than 1 mm. The designated term “micro” comes from the manufacturing techniques originally used for microelectronic fabrication.

The designated term “micro” comes from the manufacturing techniques originally used for microelectronic fabrication. Due to their ability to manufacture micro features consistently, it is now possible to integrate sensors and actuators into these devices. Through the integration of multi functions and a reduction in size, micro reactors have become highly efficient and cost effective.Micro reactors are frequently manufactured by various techniques such as “Deep Reactive Ion etching” (DRIE); the so-called “Lithographie, Galvanoformung, Abformtechnik” (LIGA) technology; micro-electro-discharge machining (µEDT); rapid prototyping; and micro- injection moulding, etc. Irrespective of the particular manufacturing technique, the basic construction of micro reactors consists of elements such as, inlet ports, outlet ports and micro channels, which define pathways for fluid flow.

In case of conventional reactors, channels are usually circular in cross section as they can withstand high pressure, but for micro reactors the cross section of the channels can be either circular or square.The flow of fluids in a duct is measured by the Reynolds number, “NRe”, which isdimensionless and is the ratio of inertial forces (Vρ) to viscous forces (µ/L) in a fluid:NRe = (Vρ L)/µ(1)where V is the mean fluid velocity (m/s), ρ is the density of fluid (Kg/m3), and µ is the dynamic viscosity (Pa.s).Equation (1) can be modified for flow through a circular pipe:NRe = (Vρ L)/µ = VL /ν = QD / νAWhere ν is the kinematic viscosity, ν =µ/ρ (m2/s), Q is the volumetric flow rate (m3/s), A is the cross section area of the pipe (m2), and D is the Diameter (M).For non-circular channels, which are described by major and minor dimensions, the “hydraulic diameter, Dh” becomes applicable.

The hydraulic diameter is defined as four times cross sectional area of the flow path, divided by the wetted perimeter:Dh= 4A/Pwhere A is the cross section area (m2) and P is the wetted perimeter (m).In theory, the Reynolds number for the laminar flow regime (NRe < 2300) is where viscous forces dominate the flow behaviour and the fluid experiences shear losses. However for NRe >2300, a turbulent flow regime exists, which is dominated by inertial forces which tend to produce eddies and rapid fluctuation.Fluid flow in micro reactors occurs mostly in the laminar regime because of the low accumulated volume of micro channels. For the design of reactors, NRe is calculated based on flow rate of the fluids; the aspect ratio of the channels; and the physical properties of the fluids. Loss of energy while flowing through these channels is defined by a friction factor “f” which is a function of NRe.

Loss of energy while flowing through these channels is defined by a friction factor “f” which is a function of NRe. Referring to the Darcy friction factor, the magnitude of friction is inversely proportional to NRe for the laminar regime. However for turbulent flow, the friction factor depends upon the ratio of surface roughness (ε) to the flow diameter (D).Consequently for laminar flow (NRe < 2300) it is apparent that the friction factor does not depend upon the flow diameter. Similarly, for NRe > 2300 (i.e. turbulent region) the friction losses in the fluid depend on the surface roughness. As micro reactors have low accumulated volumes, we can assume that micro reactors are operated in laminar regimes. However, various investigations of fluid flow in micro channels have reported deviations from the theoretical behaviour [1].

However, various investigations of fluid flow in micro channels have reported deviations from the theoretical behaviour [1]. The transition of laminar to turbulent flow has been reported to occur as early as NRe < 800 and completely changes into the turbulent regime starting at NRe in the range of 1000 - 1500.In this case, it is important to notice that the surface roughness in the micro channels is significant with respect to the flow diameter. Furthermore, occurrence of the turbulent flow regime at relatively low values of NRe results in further losses due to formation of eddies [2]. Behaviour of fluids at lower NRe therefore becomes highly complex and the experimental results show deviation from the theoretically expected values.(2)(3)Because of these observations, we are stimulated to investigate possible solutions to this problem.

Here we propose that modifying the manufacturing process can bridge the gap between the theoretical and experimental performance of the micro reactors. In this work, Digital Light Processing (DLP) based manufacturing has been exploited. DLP is based on the Texas Instruments digital micro mirror (DMD). DMD is a micro electromechanical (MEMS) device containing a matrix of micro mirrors. Each one can be individually used as a reflector to define an image pixel and to achieve high resolution projection for stereo lithographic fabrication.The surface roughness of parts created using DLP techniques varies greatly depending on how the surface is orientated relative to the build direction. The top and bottom surfaces, which are parallel to the glass substrate, can be significantly smoother than faces perpendicular to the substrate; which often suffer from surface irregularities caused by the very nature of building components in layers.

If surfaces are angled relative to the substrate than the problem can exacerbate by a phenomenon known as the ‘stair step’ effect. In order to achieve high fidelity three dimensional builds of micro reactor devices, a number of software tools have been developed which recognise features such as flat planes and changes in contour profile of the STL CAD file and adapts the position and thickness of layers accordingly. For this work to be successful it is important that the developed slicer has at the very least a comparable resolution to that of the DLP system.Software DevelopmentTo create a part using the DMD masking technique an STL file which describes the 3D solid model is firstly mathematically sliced and each slice converted into a 2D greyscale image mask (Bitmap image). The Perfactory software suite generates then assembles all the layer masks, along with an XML document containing system and build parameter settings, into a convenient JOB file (a disguised ZIP file).

The pre-processed JOB file can then be transferred into the Perfactory system for construction. The software facilitates the generation of slice data with minimal control of system parameters and with no control over layer mask generation. From a user perspective the Perfactory software is simple to use. However, the challenges set out in the work discussed in this paper requires a greater understanding of how system parameters and image masks influence hardware performance and part quality; with a view to improving accuracy, resolution and surface roughness of fine features. Previous work at Liverpool has shown that reducing the layer thickness has a significant effect on increasing DLP surface quality [3]. However, this work was limited in layer thickness range between 25 and 100 microns. This range was constrained by the Perfactory software.

However, this work was limited in layer thickness range between 25 and 100 microns. This range was constrained by the Perfactory software. The software developed in this work will enable experiments to expand on previous work by constructing test pieces with a user defined value of layer thickness and with building parameters adjusted accordingly. This section describes the development of software tools to create a JOB file; including adaptive STL slicing (FS Slicer), image mask generation and the construction of the parameter settings file.The slicing algorithm reported in this work has evolved from a machining algorithm based on CAM-style slicing with a defined tool shape. This code (written in C++ and Python) has been discussed before but some recent modifications give reason for further discussion here [4]. Conventional CAM requires the specification of a tool geometry; a cylindrical shape with atorus, spherical or flat tool tip.

Conventional CAM requires the specification of a tool geometry; a cylindrical shape with atorus, spherical or flat tool tip. The FreeSteel slicer (FS-Slicer) divides this tool shape into several parts: a tip (sphere, torus, or flat disk), a shaft (cylindrical or conical) and optionally any number of cylindrical or conical holder shapes. The shaft diameter is not restricted to the tip diameter; a shaft smaller than the tip, for example, can be used to model lollipop cutters. To start, all triangle objects contained in the STL file that describe the surface geometry are read into a single array. Each triangle object has three pointers to edge objects and triangle normal.

To start, all triangle objects contained in the STL file that describe the surface geometry are read into a single array. Each triangle object has three pointers to edge objects and triangle normal. Edge objects point to two end points and up to two adjacent triangles (it is then possible when looping over all triangles to avoid using the same edge twice or when looping over edges to consider the same triangle multiple times (maximum of 3).The slicing algorithm is then given a z height at which contour data points are required (See Figure 1a). The algorithm then builds a weave (grid) at the given z height which is constrained by an area slightly larger than the bounding box of the cross section of interest (see Figure 1b). Contours, or tool paths, are then calculated for the tip centre by passing the tool along the weave lines. When the tool shape touches the contour of the STL the contour point is recorded.

Contours, or tool paths, are then calculated for the tip centre by passing the tool along the weave lines. When the tool shape touches the contour of the STL the contour point is recorded. Triangle objects are grouped into containers so that when calculating a cutter position only the triangles that belong to the containers covered by the cutter shape need considering (prevents the need for looping over all triangles).Figure 1: (a) location of slice height in z direction, (b) slice contour and weave and (c)calculation of contour data with tool tip.When a shaft height is specified the algorithm will continually ‘look up’ for triangles above the slice height that would otherwise interfere with the tool shaft; preventing any contour data collection of undercuts (This is a desired feature of 3 axis tool paths).

Hence, for the purpose of slicing for use with SFF technologies, shaft height needs to be set to a value equal to or lower than the value of the layer thickness.The resulting contours have to then be processed through two steps: (1) Unwanted contours "inside" the volume modelled by the STL have to be filtered out (the algorithm would pass the tool along both sides of the contour because any triangles above the tool tip will not interfere with this tip) and (2) an inward offset equal to the radius of the tool tip has to be applied because contours are referenced to the centre of the tool (see Figure 1c). As the tool marches around the weave the algorithm also records contour type according to the right hand rule ofvector orientation i.e. whether the contour describes a core or cavity.

As the tool marches around the weave the algorithm also records contour type according to the right hand rule ofvector orientation i.e. whether the contour describes a core or cavity. Once the tool has passed over the entire weave, a new z height is selected and the process repeats.The resolution of the contour data is governed principally by the resolution of the weave and this is defined by the user. The size of the tool tip determines how efficiently the tool can navigate small features and internal corners. In this work weave resolution was set to 0.01mm and a flat disk for the tool was selected with a radius of 0.05mm. These parameters were chosen based on a trade-off between speed of slicing and resolution; maintaining a JOB file generation time equally comparable to the Perfactory software suite. It is also worthy of note here that a larger tool tip can be impervious to breaks in STL triangulations.

It is also worthy of note here that a larger tool tip can be impervious to breaks in STL triangulations. It is quite common to use propriety software to fix holes in STL files before they are sliced. Using a tool tip with a defined area prevents the cutter from ‘falling’ through holes in the STL geometry and sustains contour point calculations.There are also a number of additional software features that have been utilised in this work to improve build accuracy/resolution and/or reduce build time. The first is a flat planes algorithm that detects the z height position of all horizontal surfaces (see Figure 2a). This is done by looping over all triangle objects and comparing the normal’s (which are cached) with a user defined slope parameter. This is done to ensure contour data is collected close to these regions. It should be noted that a slice cannot be taken exactly at the position of a flat plane because the tool will be approaching the triangles from side on.

It should be noted that a slice cannot be taken exactly at the position of a flat plane because the tool will be approaching the triangles from side on. The second is an adaptive slicing algorithm. The algorithm works in one of two ways. Either, layer thickness value can be defined for a given z height range or the layer thickness values can be automatically modified with changes in the geometric profile of the STL file (see Figure 2b). The latter works by calculating the largest step value between the current layer zhi and the layer below zlo (see Figure 2c). If the step change is greater than a user defined threshold value then a second slice is taken half way between zhi and zlo and the calculations repeat. The iterations continue until the maximum step change equals or falls below the threshold value.

The iterations continue until the maximum step change equals or falls below the threshold value. All contour data for all layers is stored within a single dictionary where keys (z height and contour type) point to the data point arrays.Once all contour data has been collected the contour point arrays are passed into an imaging function to create the image masks for each layer. The first step is to scale the contour data according to the resolution of the DMD processing chip and the size of the projected area on the Perfactory build platform (see Equipment and Experimental Methods Section). A PNG image equal in resolution to the DMD chip is then created (1400 x 1050 pixels) and all pixels are initialised to black (0,0,0). The scaled contour points for each layer are then used to construct a series of filled polygons, the colour of which is determined by contour type in each layer (core = black, cavity = white).

The scaled contour points for each layer are then used to construct a series of filled polygons, the colour of which is determined by contour type in each layer (core = black, cavity = white). Since each polygon is drawn on the image in layers it is important that the larger contours are drawn first. Each sliced layer is centralised on the PNG image. Hence, the software currently has no capability to nest multiple parts in one build operation.Image masks created using the FS-Slicer software were comparable with image masks created using the Perfactory software suite. However, one noticeable difference with the FS- Slicer generated images was the absence of a border, one pixel in width, that surrounded each layer silhouette (see Figure 3).

However, one noticeable difference with the FS- Slicer generated images was the absence of a border, one pixel in width, that surrounded each layer silhouette (see Figure 3). However, the existence of the border in Perfactory generated layer masks was not prevalent and its greyscale intensity (ranged between values of 20 and 200)Figure 2: (a) detection of flat planes, (b) adaptive slicing and (c) calculation of sliceheight in adaptive slicer.changed in a perceived random manner when viewing a number of different image masks. This suggested on first inspection that the border may not be intentional; perhaps a consequence of some image scaling technique that implemented anti-aliasing methods. For this reason no attempt was made to recreate a border in image masks produced using the FS-Slicer software.

For this reason no attempt was made to recreate a border in image masks produced using the FS-Slicer software. Further thought is given to this phenomena in the Discussion and Conclusions Sections of this paper.To engineer a fair comparison test it was also important to be aware of the Enhanced Resolution Mode (ERM) implemented by the Perfactory system to improve part resolution and surface roughness; the mechanism by which improvements are sort was observed but not fully understood by the authors at the time of writing this paper. When ERM mode is active each sliced layer is processed twice. Each repeated layer is digitally offset from the first by shifting the image by one pixel in both x and y directions (observed by examining consecutive image masks); an analogy would be building a brick wall where bricks in adjacent layers are staggered. It is also worth noting here that the border phenomenon discussed previously often has a higher intensity in the second image mask.

It is also worth noting here that the border phenomenon discussed previously often has a higher intensity in the second image mask. The ERM methodology was implemented in the FS-Slicer software in this work.Once all image masks had been generated the FS-Slicer software saved each mask into a JOB file along with an XML file that contains all machine and layer parameter settings. The parameter settings chosen for the experiments discussed in this paper are discussed in Equipment and Experimental Methods Section.Figure 3: (a) Perfactory Image mask showing greyscale border and (b) FS Slicer Imagemasks showing no border.Equipment and Experimental MethodsFabrication of test specimens containing square and circular micro channels was done using a Perfactory Mini DLP apparatus supplied by Envisiontec, Germany.

The apparatus utilised a DMD masking technique for creating a successive number of layer masks (bitmap images) to define repeated layers of a controlled geometry by selectively photo-polymerising a liquid resin layer by layer. Parts realised in this way can be thought of as containing a large number of elemental voxels (volumetric pixels). However, the resolution of each voxel is anisotropic. This is because resolution is principally controlled in the x-y plane by the density of mirrors on the DMD chip and the focussing optics. In the z direction, resolution is governed by the stepper motor controlling the build height of the platform. In this work, projected light from a DMD chip, with a mirror density of 1400 x 1050, was focused using the Perfactory SXGA+ 85mm optics onto a build platform of size 27.77mm x 20.83mm. This gave a voxel resolution of 20 microns in the x-y plane. The resolution in z is set by the layer thickness. The stated accuracy of the system is ±7.5 microns.

This gave a voxel resolution of 20 microns in the x-y plane. The resolution in z is set by the layer thickness. The stated accuracy of the system is ±7.5 microns. These values do not take into account any material volumetric changes during polymerisation. An incandescent bulb is used as the light source. This emits a spectrum which is filtered to appear green (approximately 510nm) at the build surface. The manufacturer regards the specific wavelength to be proprietary knowledge and therefore further information is not disclosed. Greyscale can also be implemented in image masks which causes the mirrors on the DMD chip to rapidly fluctuate between their on and off states giving reduced exposure per unit time.The photopolymer used in this study was an acrylic resin; Pentaerythritottri/tetra-acrylate (CAS 4986-89-4) + 1,1,1 Tri- hydroxymethylpropyl - triacrylate (CAS 15625-89-5, 3524-68-3).

It is otherwise known as PMMA and is supplied by Envisiontec using the pseudonym R5 – a general purpose building material. The machine parameter settings as recommended for this material were used in this work. The only exception was the time of exposure (T-Exp) which was interpolated from Perfactory recommended settings using a power function for 25, 50 and 100 micron layers (see Figure 4). Gold coating of the PMMA components was done by evaporation using an Edwards S150 coating system.

Gold coating of the PMMA components was done by evaporation using an Edwards S150 coating system. Surface roughness was evaluated using a Veeco Wyko NT3310.Figure 4: Calculation of Exposure time interpolated from Perfactory recommendedvalues.Figure5: (a) Square / Circle channel block and (b) Square / Circle channel block showingvariable layer thickness.A series of cubes of size 10mm3, each differing in layer thickness (15, 20, 25, 50 and 100 microns) and by the method in which image masks were generated (Perfactory with ERM on and Off and FS Slicer with ERM on and off), were chosen to investigate surface roughness. ERM is discussed in greater detail in the Software Development section of this paper. Average surface roughness values were measured on faces perpendicular to the build direction.

ERM is discussed in greater detail in the Software Development section of this paper. Average surface roughness values were measured on faces perpendicular to the build direction. A second series of blocks, of size 24x6x8mm, were also constructed which contained a number of square and circular channels which varied in square side length / diameter from 0.1mm to 1.6mm (see Figure 5). The channel size and shape were representative of typical micro reactor channels and were built to investigate shape integrity, accuracy and surface roughness. To speed up the build process the layer thickness remote from the channels was set at 100 microns and was reduced to values of 15 and 20 microns in regions local to the channels (see Figure 5b). Building the test specimens in this way reduced the time of build by several hours.

Building the test specimens in this way reduced the time of build by several hours. Post manufacture cleaning of all parts was undertaken for 15 minutes in an isopropanol bath before secondary curing took place in air for 6 hours.Results and DiscussionAn Envisiontec Perfactory DLP system has been used to successfully construct test specimens to investigate surface roughness and build accuracy for square and circular cross section micro channels. Image masks which contain the silhouette of the cross section of each sliced layer was created using software tools shipped with the Perfactory system (PF-Slicer) and software tools written at the University of Liverpool (FS-Slicer). Dimensional analysis on the constructed test samples (FS samples and PF samples) verified the relative accuracy of the two methods for preparing JOB files.The results from surface roughness experiments are summarised in Figure 6.

Cubes of size 10mm3 were constructed using experimental conditions discussed in the previous section. As expected, for vertical walls parallel to the build direction, the surface roughness was generally improved with ERM mode active and for a smaller layer thickness. Previous work at Liverpool suggested that the combination of more pronounced voxels and deeper voids caused increased surface roughness when ERM is not active. These differences are likely caused by changes in the polymerisation process and the micro shift in light location as the second exposure occurs [3]. There were however some notable new observations that have fuelled a revised philosophy. Firstly, the surface roughness of vertical walls in parts produced using the FS-Slicer was less dependent on ERM mode. This is particularly noticeable at a 100 micron layer thickness where ERM off gave a significantly better surface finish.

This is particularly noticeable at a 100 micron layer thickness where ERM off gave a significantly better surface finish. Furthermore, for every cube built, the surface roughness was smoother in FS samples than equivalent PF samples. The absence of a greyscale border around image masks created using the FS-Slicer is the likely cause. This suggests that the reduced exposure intensity per unit time at part edges, which gives rise to less pronounced voxels at these locations, actually increases surface roughness in vertical walls. This may also account for the greatest mismatch in surface roughness between ERM modes and slicing techniques at larger values of layer thickness. The micro shift in light intensity between consecutive layers when ERM mode is active may also exacerbate the problem.

The micro shift in light intensity between consecutive layers when ERM mode is active may also exacerbate the problem. Although the latter remark is speculative at this time.Figure 6: Root mean squared (Ra) surface roughness values for vertical walls built using DLP.Figure 6 also suggests that no benefits can be gained when attempting to improve surface roughness by reducing the layer thickness below 25 microns. Moreover, with further reductions in layer thickness, from 20 to 15 microns, the surface roughness appeared to deteriorate further. Two possible reasons have been identified. Firstly, the mechanical positional accuracy of the build platform (±7.5 microns) becomes more significant as the layer thickness reduces. Secondly, the calculated values of exposure time were interpolated from only three data points which is likely to give inaccurate results.

Secondly, the calculated values of exposure time were interpolated from only three data points which is likely to give inaccurate results. This may also explain surface discontinuities observed on part surfaces which are a likely cause of residual stresses from overexposure of thin layers.Figure 7 and Figure 8 show microscope images of typical examples of square and circular channels built using the DLP apparatus respectively. In accordance with the discussion above, the square cross section channels exhibited improved shape integrity when built with image masks generated using the FS-Slicer. In general, the square channels displayed a uniform cross section throughout their vertical wall height which improved with reducing layer thickness.

In general, the square channels displayed a uniform cross section throughout their vertical wall height which improved with reducing layer thickness. This included increased squareness in channel corners and fewer surface imperfections when visual comparisons were made with PF created samples.In contrast, the circumference of circular channels were found to be highly irregular and the perimeter contour extremely pixelated. To put this observation into context, PF samples constructed with channels with the largest of diameters (>0.6mm) and with the smallest layer thickness (25 microns) could only be considered at best as “near to circular”. In comparison, circular channels constructed using the FS-Slicer software were significantly worse, demonstrating a total breakdown in the integrity of the perimeter profile for all values of layer thickness investigated (see Figure 8).

It was also speculated that the image shift associated with an active ERM mode, causing a micro shift in light intensity, aided smoothing.Conclusions and Future WorkThe absence of a one pixel wide grey scale border around the perimeter of layer masks (for both ERM modes) has been shown to give improved surface roughness in vertical walls built using DLP. Further improvements could also be made by reducing layer thickness although the selection of an appropriate exposure time becomes dominant at very small values (>25 microns). Circular channels generally exhibited pixelated perimeter profiles which only started to show good shape conformity when the ratio of channel diameter to layer thickness exceeded a value of 24 (there is scope for refinement of this value with more experimental data). This value was reached using a combination of an active ERM and a grey scale border.

This value was reached using a combination of an active ERM and a grey scale border. At the time of writing this paper the individual role played by these two enhancement techniques was not clear or which, if any, were the more dominant. The apparent arbitrary nature to the level of grey in the border regions compounded this lack of understanding and therefore gives rise to an essential area of further research. What is clear from the research is the need to use ERM and a grey scale border to improve fidelity of inclined and profiled surfaces – features which are inherent in micro reactors. Because circular micro channels were not faithfully reproducible in this work using DLP, investigations of fluid flow behaviour through them was not justifiable. The authors however believe there is opportunity to improve shape reproduction in DLP through modifications to the image masks and processing method.

The authors however believe there is opportunity to improve shape reproduction in DLP through modifications to the image masks and processing method. This solicits the question for the need of a ‘rich adaptive slicer’ in which layer masks, layer thickness and ERM are actively changed with changes in contour profile of the STL model.AcknowledgementsThis work has been supported by grants from the UK Engineering and Physical Sciences Research Council (EPSRC EP/D064856/1). In particular, we thank Martin Dunschen and Julian Todd from FreeSteel, UK, for their support in developing the STL adaptive slicing algorithms.References[1] Mohiuddin Mala, G. and Li, D. (1999). Flow characteristics of water in microtubes.[2][3][4]International Journal of Heat and Fluid Flow. 20(2): p. 142-148. Perry, R.H. (1973). Perry's Chemical engineer's handbook. McGraw-Hill chemical engineering series, 5th Edition. Hauser, C., Clare, A.T., Taylor, S., Chalker, P.R., Sutcliffe, C.J., Brkic, B.

STOMISED RAPID MANUFACTURED PARTS: TECHNOLOGY AND CASE STUDIES FROM THE CUSTOM-FIT PROJECT J. Jones, C.L.Jones* & D. I. Wimpenny Department of Engineering, De Montfort University, Leicester, England, LE1 5XY Delcam Plc, Birmingham, England B10 0JBAbstractThe design and manufacture of individually customised products is generally restricted to bespoke clothing or footwear for very wealthy customers. The aim of the Custom-Fit project was to develop a fast, flexible and economically viable route for the manufacture of individually customised parts. These products not only provide improved comfort levels but also provide better functional performance, including enhanced safety for the user. This 4.5 year, European Commission subsidised €16 million project, supported by the EU, involving 30 partners across the breadth of the Europe finished in early 2009.

This 4.5 year, European Commission subsidised €16 million project, supported by the EU, involving 30 partners across the breadth of the Europe finished in early 2009. This paper will showcase the technology developed: CAD packages which automate the design process and three new rapid manufacturing methods. It will also include case studies on a range of customised products, including customised Motorcycles helmets. The case studies not only demonstrate the performance benefits of individual customisation but also show the potential for new approaches to product design. More information at www.Custom-Fit.org1. IntroductionCustomisation has traditionally been the preserve of the rich or medically unfit, primarilydue to the cost of customized products which are often in the order of several thousands of dollars.

The driving force behind the European Custom-Fit project (2004 to 2009) was to deliver customised products at an affordable price and thus embody the concept of mass customization as proposed by Tseng and Jiao (2001) “to produce goods and services to meet individual customer's needs with near mass production efficiency”. To achieve this objective the project examined and proposed solutions across a wide spectrum of customization challenges including: business implications, technical capabilities (data capture, automated design, rapid manufacturing) and regulatory issues. The findings of the research were validated within the context of six case studies on products for use on or within the body namely: helmets, toy vehicle, motorcycle seats, knee implants (tibial plates), mandibular implants, and transfemoral prosthetic sockets.Below is a list of the customisation challenges identified prior to and during the earlystages of the Custom-Fit project.

The list was compiled from several sources including Wimpenny et al. (2000), Hague et al. (2003), Feenstra et al. (2003), coupled with the experience of the consortium members. The challenges of customisation include:Accurately quantifying a market for customisation  Cost-effectively collecting personal data  Cost-effectively modifying a product design to a personalised shape  Verifying the new design is fit for purpose and safe to use in a non-destructive way  Reliably producing parts  Delivering ALM (Additive Layer Manufacture) products with superior material properties & performanceProducing parts with better surface finishManufacturing parts faster  Certifying that a component is compliant with appropriate standards and regulationsThis paper first explores the business implications of customised products establishing evidence of a potential market value. Next, the technical aspect of the work will be considered and finally the regulatory issues will be reviewed.2.

Next, the technical aspect of the work will be considered and finally the regulatory issues will be reviewed.2. The Challenges of Customisation to Business The difficulties of successfully building a business around mass customised products arenot insignificant. One of the core difficulties is quantifying the added value of customising. Compounding this challenge is the fact that often much of the potential market has never experienced a customised product. Therefore the challenge becomes accurately anticipating the perceived value a customer places on a customised product without any tangible way for them to evaluate it. The need to measure intangibles, including the competitive advantage of customising, customer trust in a brand, and belief in an innovation has been highlighted by researchers, but to date there is no definitive predictor. (Helms et al., 2008; Dewan et al. 2000).

(Helms et al., 2008; Dewan et al. 2000). Even companies accustomed to customising can struggle to anticipate demand accurately (Piller, 2008).The Custom-Fit project undertook market research in the form of an online survey toestablish the market demand for customised motorcycle seats. Ong et al. (2008) have published the results of the survey which summarize the responses of over 3000 individuals, mostly men (almost 98%). The survey found that less than 8% of motorcycle drivers share their bike with other riders. 92% of riders experience some discomfort from their seats when travelling long distances. Over 80% of responses favoured the idea of a customised motorcycle seat and were willing to pay more for it (66% up to 250 Euros and 30 percent up to 500 Euros).Based on the results of the above Custom-Fit survey (Ong et al. 2008) one can extrapolate the potential European market value.

2008) one can extrapolate the potential European market value. If eight million (approximately two-thirds of the European motorcycle riders) bought a customised seat at 250 Euros the market would be valued at 2,000,000,000 Euros. Even taking a very conservative view of the survey results and suggesting that perhaps only one percent of the riders would actually buy a customised seat then the market value is still 2,000,000 Euros. While these numbers are wildly speculative, the interest demonstrated in the survey is clear and suggests that there is a market for customised motorcycle seats.3. The Custom-Fit ApproachThe Custom-Fit project approached customisation from a simple paradigm: Use customer data to automate the design and manufacturing chain of customised parts for use on or inside the body (Jones and Wimpenny 2008).

The mass customising of garments illustrates the effectiveness of this philosophy where individual measurement information can be integrated to adapt generic patterns which drive the manufacturing process (Helms, et al. 2008). While this approach is well developed and automated in applications where designs are typically two- dimensional (2D) such as the garment industry, the adoption of this philosophy for three- dimensional (3D) products has been limited primarily due to the need for human intervention at the data capture and design stages. Implicit in the scope of this work was to develop a set of software tools to automate a generic 3D design chain competent across a wide variety of products.

Implicit in the scope of this work was to develop a set of software tools to automate a generic 3D design chain competent across a wide variety of products. Each step in the design and manufacturing chain will now be considered as shown in Figure 1.Data CaptureNSFAutomated DesignSRIAdditive ManufacturePartMarketGeometrical Data  Non-geometrical data  Subjective user preferencesMorph a standard CAD model to match a specific customer  Internal topography  Virtual validation andSlice the customized modelRapid ManufactureIssue a certificate of conformanceCustomer Sign OffverificationFigure 1 - The Custom-Fit Approach3.1 Data CaptureThe backbone of customer-centric design isto effectively collect as much relevant data as possible from the customer. In the last decade the advances in 3D shape measurement and capture have played an increasingly important role in a variety of fields including manufacturing and medical sciences (Zhang, 2009).

In the last decade the advances in 3D shape measurement and capture have played an increasingly important role in a variety of fields including manufacturing and medical sciences (Zhang, 2009). 3D non-contact laser scanning was adopted as a standard approach for all products except the implants (VITUS scanner solution by Human Solutions www.human- solutions.com) to quickly (less than 10 seconds) and efficiently capture the shape of customers for the design of customised products. In addition to the physical shape of customers the helmet and motorcycle seat products required pressure distribution information which was captured using an array of pressure sensors integrated into a flexible “blanket” (X3 PX100:18.18.01 seat sensor by XSENSOR Technology www.xsensor.com). The last type of data collected was subjective.

The last type of data collected was subjective. Through a series of questionnaires (developed by the Instituto de Biomecánica de Valencia www.ibv.org) the customer identifies areas on the body of pain or discomfort and assigns a magnitude as shown in Figure 2. This data can then be used within the design process and also for product evaluation. The final need addressed under the data capture umbrella is the ability to package it in a file format that can be readily used by other software tools in an automated way. The shortcomings of point cloud, meshes, and other scan formats has been identified by various researchers (Wand, et al., 2008; Stroud and Xirouchakis, 2000). Additionally, the scope of applications required a more universal container file capable of holding customer details, 3D mesh data, DICOM data, pressure maps, and customer preferences. The consortium found the need to establish a new file format for exchanging data between software applications.

The consortium found the need to establish a new file format for exchanging data between software applications. The new file format is called the Neutral Scanning Format (NSF) and is a container capable of holding a wide variety of data types in XML format.Figure 2 - Comfort Questionnaire3.2 Automated DesignOnce the NSF file was filled with customer data it the automated design process begins.Custom-Fit built its design automation approach on the concept of “reuse” as originally proposed by Prieto-Diaz (1993) and later reinforced by Tseng and Jiao (1997). Most of the design is “reused” or recycled for each product and only the features that are in contact with the customer during use are customised. Therefore the first step of the design process was to select a base design for further customisation. The 3D scan data was interrogated to establish a match with the closest standard product model available.

The 3D scan data was interrogated to establish a match with the closest standard product model available. Next the scan was oriented within the same coordinate system as the standard CAD model so that there was a small gap or overlap between them. The standard product was then morphed or a Boolean operation was performed to conform exactly to the scanned shape.Having achieved a customised form, the next focus is on enhancing the functionalitythrough modification of the mechanical properties of the product. This is achieved through functional grading of the materials used and/or employing internal lattice structures. The knee implant and prosthesis used the former and the mandibular implant, helmet and motorcycle seat used the latter.

The knee implant and prosthesis used the former and the mandibular implant, helmet and motorcycle seat used the latter. Finite element analysis and a design approach were used to establish how the materials or lattice structures should be distributed.Once the internal and external customisation of product is completed the performanceunder service conditions needs to be checked. The consistency of approach and reliability of the output is one of the criterions that distinguish mass customisation from craft production (Piller, 2004). For that reason each design created by automated design tools is tested with FEA or FEA established bandwidths prior to manufacture. In this way the high variety of designs can cost- effectively be validated.3.3 Additive ManufactureOnce a validated design was completed it was then advanced on to the manufacturingstage. As is typical with additive manufacture the models were sliced in preparation for a build.

As is typical with additive manufacture the models were sliced in preparation for a build. Where graded materials had been assigned to a design STL files could not hold all of the relevant information so a slicing methodology known as a Slice Raster Interface (SRI) format was developed so that assigned material distributions could be translated into a stack of multi- coloured bitmap slices where each colour corresponded to a different build material.Throughout the project new designs were validated to extent possible using existingRP/RM (Rapid prototyping/manufacturing) systems, however as the project drew toward conclusion and the capability of the three new RM systems increased the production of parts shifted toward the new systems. An explanation of the new systems follows in the next section.3.4 MarketCareful consideration must be given to the development of a robust methodology for testing of customised products to ensure that they comply with safe performance criteria.

For example According to UN regulations no.22, a motorcycle helmet has to pass a series of 12 tests before it is considered “suitable”. Fortunately, in many cases customised products offer higher levels of user safety than mass produced products. For example, a poorly fitting motorcycle helmets can give discomfort leading to higher levels of rider fatigue and reduced concentration. Moreover, a customised helmet should ensure more effective distribution of impact stresses across the rider’sskull. The precise testing methodology depends to some extent on the nature of the product but may be based on a combination of empirical tests and computational simulation (this is discussed in more detail in section 7).In addition to meeting safety performance requirements it is important that some form of effective customer sign-off is included in the process. The customer sign-off process provides vital data on customer satisfaction levels to enable the overall process to be fine-tuned.

The customer sign-off process provides vital data on customer satisfaction levels to enable the overall process to be fine-tuned. Moreover, forming a strong link with the customer is a critical part an effective business model for customised products. From the outset close communication with the customer is required to record their requirements/desires for the product and at the same time manage their expectations. Compromises often have to be made to ensure the product meets safety requirements can be manufactured effectively, as well as providing enhanced customer comfort levels. Another added complication is that some customised products may only demonstrate their true value in terms of increased comfort after several hours of use.4. New Manufacturing Technologies As demonstrated by Wimpenny et al. (2000) rapid prototyping techniques need to be more reliably deliver better accuracy, speed, and surface finish.

New Manufacturing Technologies As demonstrated by Wimpenny et al. (2000) rapid prototyping techniques need to be more reliably deliver better accuracy, speed, and surface finish. Piller (2008) summarized the current limitations of ALM and identified the need to scale up production to enable mass manufacturing through rapid prototyping. (Piller, F. 2008). Three new processes were developed in the Custom-Fit project. The team at TNO (Eindhoven,Netherlands) were responsible for managing the development of the High Viscosity Inkjet printing (HVIJ) process. This approach is based on continuous inkjet technology but the unit is configured to enable viscous materials to be deposited. A test rig was developed to enable a mandibular implant with a complex internal lattice structure to be produced using a UV curable bioresin (see Figure 3).Figure 3 – The high viscosity inkjet printing test rigThe Metal Printing Process (MPP) was developed by the team at Sintef (TrondheimNorway).

In this system a layer of metal powder is selectively deposited and then fused under high pressure within a heated chamber (see Figure 4). As opposed to laser/electron beam sintering the MPP process is effectively solid state sintering which allows very high integrity metal parts to be produced, including functional grading of several metals without appreciable alloying (Boivie 2006). In the Custom-Fit project MPP was developed for manufacturing the tibial implant.Figure 4 - The MPP systemFigure 5 – SLP test rig at MTT LtdIn the 3rd system developed, electrophotography (laser printing) is used deposit a precisepatterned layer of a polymeric toner which is then fused using a radiant IR heat source to produce a solid thermoplastic object. In addition to relatively high production speeds the Plastic Printing Process, now known as Selective Laser Printing (SLP), has the potential to produce complex functionally graded materials using several thermoplastics in a single part.

The SLP process was developed by DeMontfort University (Leicester, UK), MTT Technologies Ltd(Stone UK) and CTG GmbH (Alsdorf, Germany). A test rig was constructed to produce the seat, helmet and prosthesis demonstration parts in the Custom-Fit project (see Figure 5).5.

A test rig was constructed to produce the seat, helmet and prosthesis demonstration parts in the Custom-Fit project (see Figure 5).5. Case StudiesDuring the Custom-Fit project 6 case studies were undertaken to assess/demonstrate the potential of individually customised products;Motorcycle helmet  Motorcycle seat  Transfemoral prosthesis  Tibial (knee) implant  Mandibular implant  Toy car seatAlthough all of these cases studies successfully demonstrated different aspects of the benefits and challenges of introducing customised products, only the motorcycle helmet and mandibular implant will be described in detail in this paper.5.1 Motorcycle Helmet: Bespoke motorcycle helmets have the potential to offer superiorprotection in terms of distributing crash forces over a larger area of the skull and also remaining in an ideal crash protection position on the head in the event of an accident.

Figure 6 shows various stages through the virtual creation of a bespoke helmet insert comforted to the scan of the intended customer. Step 1- In this case the customer’s head geometry was scanned using a body scanner (Human Solutions) and customer feedback was collected via a questionnaire (IBV). Step 2- That information was fed into DelCAM’s Powershape platform to morph the standard helmet liner to conform to the customers head. The liner was further functionalized for ideal stiffnessand airways for cooling by adding an internal lattice structures using Marcam’s VisCAM software(see Figures 6 & 7)..Figure 6 – morphing of standard linerFigure 7 - customised linerStep 3- The morphed design is verified using FEA software from BPO to ensure that the design customisation is prevented from compromising safety (see Figure 8). Step 4 - The liner is produced using an existing rapid prototyping process, such as laser sintering, or by the new Selective Laser Printing process.

Step 4 - The liner is produced using an existing rapid prototyping process, such as laser sintering, or by the new Selective Laser Printing process. The result is an affordable customised helmet insert which comforts to all certification standardsFigure 8 - virtual testing of customised liner5.2 Mandibular Implant: Complex maxillofacial surgery maybe required to correct congenital defects, treat victims of trauma or remove advanced head/neck tumours. Unfortunately, the current conventional approach is to undertake hard tissue reconstruction using bone harvestedfrom a second surgical site. In addition to the associated surgical trauma and risk of infection it is almost impossible to obtain a perfect geometrical match for the required tissue. The use of customised metallic (typically titanium) implants, produced to the desired shape, is now becoming a more widely accepted surgical approach.

The use of customised metallic (typically titanium) implants, produced to the desired shape, is now becoming a more widely accepted surgical approach. However, many clinicians would prefer to utilise materials, such as bioceramics and polymers which can be formulated to resorb and be replaced with natural bone over time. In the Custom-Fit project the team at AZM (Maastricht, Netherlands) headed by Jules Poukens designed a customised implant to treat a mandibular defect. CT scan (dicom) data was loaded into the NSF system which, in addition to providing a way of recording patient information in a transferable file format, also tracks how the data is manipulated at each stage of the process to provide traceability for quality assurance.The data was loaded into the maxillofacial implant design system, based on a combination of Mimics and 3Matic software, developed by the team at Materialise (Leuven, Belgium).

The external geometry of the implant was designed and then InnerSpace software developed by TNO was used to transform the solid implant into a lattice with graded porosity to enable effective osseointegration of the implant. The data was then supplied to the HVIJ printing machine for printing of the bioresin implant. The new bioresin material developed in the Custom-Fit project is still undergoing the regulatory approval process and so the new implant could not be evaluated in a patient, however, accuracy and mechanical property trials were undertaken on the implant produced (see Figure 9).Figure 9- bioresin implant inserted into model of mandible6. Certification and Regulatory IssuesOne of the major challenges with individually customised products is the need to ensure that products comply with the prevailing legislation with respect to performance. This is particularly important for products with inherent safety implications (in principle this could apply to any product).

This is particularly important for products with inherent safety implications (in principle this could apply to any product). In the case of the custom-Fit case studies the helmet and prosthesis have important mechanical performance criteria which must be met. Regulatory issues are even more onerous for medical devices which will be implanted (for example the tibial and mandibular implants). The conventional approach to testing of products used for series production is impractical for on- off products, as this usually requires multiple tests to be conducted on products to failure.

The conventional approach to testing of products used for series production is impractical for on- off products, as this usually requires multiple tests to be conducted on products to failure. To enable customised products to be introduced requires a different approach based on either virtual testing of each customised product or testing of indicative products which span the range of customisation or limit the customisation to elements of the design which do not influence the safe performance of the product (this can lead to very inefficient designs).In the Custom-Fit project, Dutch partner BPO investigated the restrictions imposed by the current statutory regulations and explored several alternative methodologies which could be applied to ensure the safe operation of individually customised products.

Trials undertaken byBPO over many years, comparing the results of FEA predictions with the tests on real products, have shown that Computer Aided Verification (CAV) methods can provide an accurate and reliable prediction of the performance of products (Posthuma and Jansens 2007). Indeed in some cases the computer predictions can prove to be more effective as they can provide more incisive information than a simple pass or fail load test on real product.

Indeed in some cases the computer predictions can prove to be more effective as they can provide more incisive information than a simple pass or fail load test on real product. Unfortunately, although optimisation of designs based on the results of FEA is common practice, to replace physical testing with CAV will require extensive changes to product legislation.Based the work undertaken in the Custom-Fit project BPO have proposed a radical new philosophy; Computer Aided Type Approval and Process Approval based on four approaches; Destining approach, Channels or bandwidth, approach, Iterative approach and Analytical approach (see Figure 10).Idea generationConcept developmentChain approvalMarket introductionSales Engineering Computer aided type approvalFigure 10- Product development process using Computer Aided Type Approval and Process Approval.The destining and channels approach as used to gain type approval.

In the destining approach the performance of a particular customised product is predicted and if it meets the performance criteria it will be produced. In the channels approach the limits of product customisation will be virtually tested and all products between these limits are assumed to be safe.The iterative and analytical approaches are used to gain chain approval. Chain approval ensures safe products by using a robust process chain in which product performance validation is integrated. In the iterative approach a rapid succession of analysis is conducted and used to refine the design until satisfactory performance is achieved.

In the iterative approach a rapid succession of analysis is conducted and used to refine the design until satisfactory performance is achieved. In the analytical approach the design of the product is driven by design rules which define relationships between performance and product geometry through a KBE (Knowledge Based Engineering) route.Of these approaches the analytical is probably the most elegant and in principle should be the most efficient, although significant investment is required to generate the design rules in the first place.7.

DiscussionAlthough significant progress has been made as a result of the Custom-Fit project, further work is still required to transform the dream of mass production of individually customised products to become a reality.Production7.1 Market for customised products; although surveys conducted during the Custom-Fit project indicate a market for some individually customised products the information gathered is very limited in scope and tends to relate to niche products or customer groups. To obtain more meaningful (widely applicable) data is somewhat difficult since the benefit of customisation is not always clear to people who have not tried it.7.2 Cost of collecting personal data; the custom-fit project highlighted some serious practical problems in the area of body scanning.Lack of 3D scanning equipment outside the medical world (except for a few pioneers like the footwear industry).Cost of employing someone to scan you.

 Cost of cleaning up scan data to provide a watertight file.  Need for the NSF file (basically an XML based container).7.3 Design customisation; despite the impressive results demonstrated to date, the design of customised products is still not fully automatic and represents a significant bottle neck and cost within the product development chain. Moreover, each product type requires its own design automation system which will be expensive to develop.7.4 Design verification/certification; In addition to the cost and time required to verify the design of each customised product, extensive changes to product legislation will be required before the CAV approach can be widely adopted.7.5 Manufacturing; Despite developments in Additive Manufacturing Technology the approach is still only viable for relatively small complex products, where a high quality surface finish is not essential.

Moreover, we are currently hampered by the limited range and cost of materials available (particularly for plastic products).7.6 Business integration; further work is urgently required to understand how customised products can be integrated within the current manufacturing and supply chain.Finally if all of the technical challenges can be overcome care must be taken to avoid “over” customising products beyond the point of diminishing returns (Dewan et al. 2000).ReferencesHelms, M.M., Ahmadi, M., Jih, W.J.K., Ettkin, L.P. (2008) “Technologies in support of masscustomization strategy: Exploring the linkages between e-commerce and knowledge management”, Computers in Industry, Vol 59, pp. 351–363Jones, J. And Wimpenny, D.I. (2008) “Digitally Customising Products for yourBody, Bike, and World”, presentation at the Time Compression Technologies 2008 RapidManufacturing Conference, Coventry, UK.Piller, F.T.

BER-PHYSICAL VUNERABILITIES IN ADDITIVE MANUFACTURING SYSTEMSL. D. Sturm, C. B. Williams, J. A. Camelio, J. White, R. ParkerDesign, Research, and Education for Additive Manufacturing Systems LaboratoryDepartment of Mechanical EngineeringVirginia Polytechnic Institute and State UniversityAbstractOne of the key advantages of additive manufacturing (AM) is its digital thread, which allows for rapid communication, iteration, and sharing of a design model and its corresponding physical representation. While this enables a more efficient design process, it also presents opportunities for cyber-attacks to impact the physical word. In this paper the authors examine potential attack vectors along the Additive Manufacturing process chain. Specifically, the effects of cyber- physical attacks, and potential means for detecting them, are explored. Based on the results of this study, recommendations are presented for preventing and detecting cyber-physical attacks on AM processes.1.

Based on the results of this study, recommendations are presented for preventing and detecting cyber-physical attacks on AM processes.1. Cyber-Physical Vulnerabilities in Manufacturing Processes1.1 Attacks on Cyber-Physical SystemsCyber-physical systems (CPS) are systems that integrate physical hardware with software systems, often with the use of a network. With the growth of the Internet of Things (IoT) the number of CPS systems on networks continues to increase [1]. Concurrently, cyber-attacks have become more prevalent in recent years, increasing in maliciousness and decreasing in visibility [1-3]. This poses a significant issue, as cyber-attacks on cyber-physical systems could result in damage to the machines themselves or the humans who interact with them.A prominent example of a cyber-physical attack was the Stuxnet worm that targeted Iranian centrifuges used for refining uranium.

In this attack the worm was able to infect the software system and affect the physical hardware, causing damage to the centrifuges. By sending false data back to the operators, Stuxnet was able to make it appear as though the centrifuges were operating correctly, while it caused them to damage themselves. The ability of Stuxnet both to cause damage to physical systems and to hide itself illustrates the ability of a cyber-physical attack to disrupt manufacturing systems and the need for physical methods of detection [4].Another example of a cyber-physical attack is the hijacking of insulin pumps. In this case a hacker is able to connect to a Bluetooth enabled insulin pump to control the dose of insulin given to the wearer. By increasing or decreasing the dosage of insulin, it is possible to cause serious injury or even death in the user.

By increasing or decreasing the dosage of insulin, it is possible to cause serious injury or even death in the user. The currently security system for these pumps is insufficient to prevent a cyber-physical attack that could have potentially lethal consequences [5].1.2 Cyber-Attacks on ManufacturingThe previously mentioned examples demonstrate the ability for cyber-attacks to cross over into the physical world. Attacks on CPS are even more alarming when considering the ever- increasing amount of networked devices that are being connected to machines in the manufacturing world. A cyber-attack on these machines could cause injury to plant workers and damage to the machine itself. More insidiously, an attack could be designed to cause a process to produce faulty parts that might find their way into end-user products [6].

More insidiously, an attack could be designed to cause a process to produce faulty parts that might find their way into end-user products [6]. For example, an attack could be designed to affect the production of a jet turbine part such that it would pass inspection but fail during operation and cause significant damage.With the rise in both the number of CPS connected to networks and in malicious cyber- attacks, there is a clear need for research to understand the vulnerabilities of cyber-physical systems. While methods exist for detecting cyber-attacks on computer systems, no such research has been done on detecting an attack from the physical parts created by the attack. As such, the authors have begun an investigation on cyber-physical vulnerabilities in manufacturing systems [7-8].1.2 Context: Cyber-Physical Vulnerabilities in Additive ManufacturingIn this paper, the authors scope their research solely on Additive Manufacturing (AM) systems.

The process chain of these networked machines has unique vulnerabilities that warrant a detailed investigation due to their ability to fabricate parts in a layer-wise fashion. For example, voids can be placed inside of a part and the material properties of internal layers can be changed without affect the exterior layers. Because of the potential damage from a cyber-physical attack, there is a need to look at AM systems to determine what vulnerabilities exist and how to prevent and mitigate the threat of cyber-attacks.An overview of the AM process chain is presented along with a description of potential vulnerabilities in Section 2. Based on this evaluation, a cyber-physical attack case study is presented in Section 3 in which the .STL file structure is attacked. The effectiveness of this attack is evaluated through part testing (Section 4). Finally, in Section 5, the results of the attack are analyzed to identify ways of preventing and mitigating future attacks.2.

Finally, in Section 5, the results of the attack are analyzed to identify ways of preventing and mitigating future attacks.2. Cyber-Vulnerabilities in the Additive Manufacturing ProcessTo be able to prevent a cyber-attack, one must first understand the vulnerabilities and weaknesses of the system. To do this, it is necessary to follow a cyber-attack through the process chain, from conception to simulated deployment. In this section, the AM process chain is examined for potential vulnerabilities to cyber-attacks.The digital nature of the additive manufacturing process chain, shown in Figure 1, provides an opportunity for a cyber-attack to cross into the physical world. There are four main steps on the process chain where an attack could take place: the CAD model, the .STL file, the toolpath file, and the physical machine itself.Figure 1.

There are four main steps on the process chain where an attack could take place: the CAD model, the .STL file, the toolpath file, and the physical machine itself.Figure 1. Additive Manufacturing Process Chain2.1 CAD ModelThe CAD model is the first step in the additive manufacturing chain and is common both to additive and subtractive manufacturing. This step in the process is the most valuable in terms of information, as it contains all of a part’s geometric data. If connected to a PLM software, the data could also include information related to simulated performance (e.g., results from finite element analysis, computational fluid dynamics, multi-physics simulations, etc.) and the associated parameters of the part’s intended use.Attacks at this point in the process may focus on stealing or corrupting files. One example of this is ACAD/Medre.A worm, which was designed to infect and steal AutoCAD drawings [8].

One example of this is ACAD/Medre.A worm, which was designed to infect and steal AutoCAD drawings [8]. Another example of this is the CryptoLocker malware, which encrypts a user’s files and then demands a ransom for the encryption key to unlock these files [9]. Because CAD files are the basis from which other drawings and files are generated, the theft or loss of these files can be costly. Furthermore, a competitor who stole these files would have access to all of the design steps that went into making the CAD model and could easily reverse engineer or modify the part.The complex nature of most CAD files makes it more difficult to directly alter the part; however an attack could be designed to do so. For example, a CAD file for a crankshaft could be altered to reduce the area of the load bearing member, resulting in premature failure. Any corruption at this phase would propagate through the entire process chain, resulting in a part that is bad from start to finish.

Any corruption at this phase would propagate through the entire process chain, resulting in a part that is bad from start to finish. However, because parts may still be edited during the CAD phase the chances of detection of an attack increase.2.2 .STL/.AMF FileThe second step in the process chain is to convert the CAD model into a .STL file, the current defacto file standard in additive manufacturing. Upon conversion, most of the model data is lost, and only the surface information of the part remains. At this stage the model is no longer represented by complex mathematical equations, but by a simple surface geometry composed of triangular elements, called facets. Each facet is defined by three vertices, specified by a set of x,y, and z coordinates Each closed set of facets comprises a shell, and multiple parts can be represented inside a single .STL file by including multiple shells.As with the CAD file, the part could be altered to reduce its performance.

By changing the coordinates of the vertices it is possible to edit the model. Additional vertices can also be added to create new features. Additive manufacturing is unique in that the interior of a part can be altered without affecting the exterior of the part, resulting in a part that looks and feels strong, but is weak on the inside.Section 3 gives more details on the potential presented in .STL attacks and on why this attack vector was chosen as a focus.. Both the .STL and .AMF file (the current ISO/ASTMstandard) are simple file formats that contain vertices of facets and so are both susceptible to the same types of attacks. The .AMF file contains more information that makes it vulnerable to a longer list of attacks than the .STL.

The .AMF file contains more information that makes it vulnerable to a longer list of attacks than the .STL. For this reason the focus will be on the .STL file, with the understanding that these techniques are also valid on the .AMF file.The CSG history of how the geometry was created is no longer present in the file, so it becomes harder to reverse engineer the information needed to create a similar part. Despite this loss of information, a theft of a .STL file is still costly as it (i) contains all of the information needed to fabricate the geometry of the part (which could result in the production of counterfeit copies) and (ii) the surface geometry data can be attacked to nefariously change part geometry.2.3 Toolpath FileUpon receiving a .STL file, each additive process converts the model into layers and generates a toolpath from these layers.

Conceptually similar to GCODE, this toolpath file contains the commands for the controllers that move the AM systems’ coordinate axes and deposition mechanisms (e.g., extrusion federate, laser power, inkjet pulses, etc.).Potential attacks on the toolpath are to place/remove material in the wrong location, to cause layers to be placed to close/far away from each other, and to damage the part/machine by driving the tool into the part/machine. Toolpath attacks offer the most freedom in what an attack can achieve because any operation the machine could normally use when creating a part can be altered in the toolpath. A toolpath file could be intercepted by a virus on the machine’s computer or when the file is sent from one computer to the machine. For example, with a wireless printer the information might be sent over a network where it could be intercepted.Theft of the toolpath file is less valuable that than that of the .STL file, since it is machine specific.

While it could still be reverse engineered or used on another machine of the same type to produce the part, it requires more work to do so than an .STL file. Attacking the model through the toolpath does allow any of the changes that can be made in the .STL file to be made to the part, however the implementation of such changes is more difficult in the toolpath.2.4 Physical MachineThe final step in the AM process chain is the machine itself. This stage of the process is vulnerable to Stuxnet type attacks, in which a worm alters the firmware of the system’s controller PC. It is noted that such a sophisticated attack vector is not needed, as most AM systems have open USB ports for maintain the machine’s software.Using this as an attack vector, the machine’s process parameters (e.g., nozzle temperature, laser intensity, etc.) could be altered to affect the material properties of the part.

For example, inkjet nozzles might be turned off at one point to prevent material from being placed, or turned on at another point to place excess material. Temperatures of extrusion nozzles could be altered to change the mechanical properties of the extruder or to clog the nozzle all together. Laser intensity could be altered to provide lower energy density to the bed to change the final part properties. Figure 2 shows configuration data for a layer being printed on a Connex 350 printer that was intercepted in plaintext using WireShark [7]. By intercepting and altering this configuration file it would be possible to remotely change the material properties by altering the UV intensity. This type of attack can be difficult to detect without some type of independent monitoring.The two limitations on this attack method. First, no model data is available to be stolen or altered. Second, any attack must be specifically targeted to each machine.

First, no model data is available to be stolen or altered. Second, any attack must be specifically targeted to each machine. The difficulty of attacking the machine can also vary significantly from one machine to another. While one machine might be particularly vulnerable to an attack, another machine could be well secured and hard to infect.http://en.wikipedia.org/wiki/File:FDM_by_Zureks.pngPrint Slice ImagesConﬁgurationSafety BypassPrint Head TemperatureFigure 2. Objet configuration data intercepted from WiFi using WireShark [7].3. Case Study: Cyber-Attacks on the .STL FileThis section will cover the development of a cyber-physical attack by examining some of the considerations that go into attacking a .STL file. The .STL is the one potential attack that does not require specific modification for each AM machine as it is an attack on the standardized file format that every AM machine takes as input.

The .STL is the one potential attack that does not require specific modification for each AM machine as it is an attack on the standardized file format that every AM machine takes as input. As STL file creation occurs at the beginning of the process chain, and is general to every AM machine, a focused attack on the file could have severe implications and could cripple an AM production line. To prevent replication of this attack, specific details of the final attack algorithm have been omitted.3.1 Types of AttacksAs described in Section 2.2, while a .STL file only has surface data in the form of a list of triangle vertices, there are several different types of file attacks that can affect the final part geometry. The effects of these attacks on a sample part geometry (an ASTM tensile test specimen) are illustrated in Figure 3.Corruption/Encryption – A traditional cyber-attack where the file is damaged or encrypted, this renders it inaccessible to the user.

A corruption or encryption attack is a straightforward attempt to damage or extort the owner of the file. By rendering the file unusable it makes it very evident that an attack has taken place. However, if backups arenot available, this type of attack can cause a lot of damage before it is detected. (Figure 3a)Scaling – The part is scaled up or down in one or more axes resulting in a changed form that may affect the fit or strength of the part. For example, a tensile test specimen that is 10% thinner may not be noticeably different to the eye, but will cause a measurable change to performance. This type of attack is easily detected in software or by measuring the completed part. (Figure 3b) Indents/Protrusions – Small protrusions or indents may be added to a part to affect the fit, surface finish, or strength of the part. For example, the inside of a printed duct or pipe might contain indents that negatively affect the flow.

For example, the inside of a printed duct or pipe might contain indents that negatively affect the flow. These attacks are visibly detectable, but may be placed in locations where they are difficult to see or measure. (Figures 3c and 3d)Vertex Movement – One or more vertices in the part is moved, resulting in a changed form that may affect the fit or strength of the part. This has the advantage of manipulating the shape of the file without changing its size. Unlike scaling, vertex manipulation allows almost any section of a part to be sized both up and down, as well as warped or altered. Measurements can detect this attack, but by altering areas that are difficult to measure, the presence of this attack can be hidden .(Figure 3e and f)Voids – Similar to an indent, a void is an area where material has been removed from the model. A void differs from an indent in that it is completely enclosed inside the existing model.

A void differs from an indent in that it is completely enclosed inside the existing model. Because voids are completely enclosed they are undetectable by dimensional measurements and may be difficult or impossible to find visually. The use of supporting material in many processes also renders the void undetectable by weighing, since the void is filled with a structurally deficient, but equivalently dense material. The presence of a void weakens the strength of a part, and if placed in a loadbearing location, may cause part failure. This attack does increase the file size slightly, but the small change is unlikely to be noticed.ABCDEFFigure 3. .STL Attacks on an ASTM dogbone. A) An unaffected dogbone B) A scaled down dogbone C) An indentation D) A protrusion E) A vertex moved inward F) A vertex moved outwardThe first four types of attacks are shared by both additive and subtractive manufacturing and can have a demonstrable effect on the part being produced [8].

The last attack, the creation of voids inside of a part is unique to additive manufacturing due to the layer based fabrication method. Because voids are difficult to detect once a part has been created and because they are an attack that is unique to AM, this specific attack type was chosen for further investigation in this study.3.2 Void Attack ConsiderationsThere are several factors that must be considered when attacking a .STL file using a void:File Size – When altering an .STL file a change in file size can may make it easier to detect an attack. Some attacks, such as scaling or moving a vertex, do not change the file size. Other attacks, such as adding voids or other features, will change the file size depending on the complexity of the feature added. It is especially concerning that a void can be added by inserting as few as four additional facets, an increase of only 200 bytes.

It is especially concerning that a void can be added by inserting as few as four additional facets, an increase of only 200 bytes. This size increase is negligible in all but the simplest files, as .STL files can easily be in the megabyte range.Location – The location of a void in a part has a significant impact on the effect of the attack on a part. A void is not a threat if it does not result in any reduction in part strength. A well place void is a concern, as it can lead to a critical part failure.Void Shape – The shape of the void affects the amount of information needed to represent it in a .STL file. The shape can be as simple as a tetrahedron – adding only four facets and 200 bytes to the file.Void Size – Void size is important for affecting the mechanical strength of the part. A void that is too small will not be fabricated, due to process resolution limits.

A void that is too small will not be fabricated, due to process resolution limits. Large voids increase the chance of part failure, but may be easier to detect.Void Number – The number of voids placed can be varied based on the desired effect. A larger void might be made less noticeable by replacing it with several smaller aligned voids. The addition of more voids will add more to the file size and may make the change more noticeable.Full Enclosure – It is essential that any void placed is fully enclosed in the original part. If a void is not fully enclosed inside of a single shell it will cause printing errors that can alert the user.Run Time – The time it takes for the attacking software to analyze and place a void is directly related to how easy it is to detect. A several millisecond delay is likely to go unnoticed by a user, while a several second delay is likely to draw attention.Silent – The software does not set off any warnings.

A several millisecond delay is likely to go unnoticed by a user, while a several second delay is likely to draw attention.Silent – The software does not set off any warnings. For example if the void contained a facet that had the normal facing in the wrong direction Netfabb would display a warning to the user alerting them that there was a problem with the file.The most dangerous void attack would be one where the void was located in a structurally important location and sized such that it was difficult to detect while still causing a critical structural failure in the part. Such an attack would run quickly, add a minimal amount of data to the part file, and would not cause any warning messages when inspecting or printing the part.3.3 Attack EmbodimentAs mentioned, it was decided that a void attack would be used for the study.

This attack waschosen for the following reasons:Uniqueness – The ability to place voids inside of a part is unique to additive manufacturing.Visibility – Unlike a protrusion, indentation, or other external change, a void cannot be physically measured once it is enclosed inside of the part. Because of this, it was desirable to determine what would or would not make an internal void visible. Most quality control processes rely on dimensional checks or visual inspection methods, such as operators or computer vision systems.

Most quality control processes rely on dimensional checks or visual inspection methods, such as operators or computer vision systems. By making the void invisible, these quality control processes can be subverted.Threat – Because voids are not readily visible when looking at a model on a computer screen, and are not measurable once the part has been made, they present a higher likelihood of passing undetected.The attack software featured three key pieces of functionality: (i) the ability to automatically determine a location for the void, (ii) the ability to ensure that the void is located completely inside the part, and (iii) the ability to automatically scale the void based on the part. In addition the voids placed did not set off warning messages in Netfabb, Objet Studio, or any of the other software we worked with the parts in.

In addition the voids placed did not set off warning messages in Netfabb, Objet Studio, or any of the other software we worked with the parts in. We developed this software in order to assess its potential impact, to gain an understanding of what the effects of such an attack were an a part, to see if it could be detected through existing quality control measure, and if not, to see what new control measures would be needed to detect it.Void PlacementIn the attack software presented there are two methods for locating voids, random placement and targeted placement. In the random placement mode, a random location inside of the part is chosen and the void is placed there. The targeted mode is a semi-intelligent method for locating the void. Because .STL files are not solid models it is difficult to attempt any type of structural analysis of the part.

The targeted mode is a semi-intelligent method for locating the void. Because .STL files are not solid models it is difficult to attempt any type of structural analysis of the part. Additionally, since the parts infected could vary significantly, and the purpose of each part is not known, it is hard to place a void in any meaningful way. To solve this problem the software uses a semi-intelligent method to place voids in locations that are likely to have stress concentrations. Areas with holes or sharp changes in curvature are often located at important locations in a part. To detect these areas in a .STL file the density of the mesh can be estimated. Similarly to FEA, more complex areas will be represented by a denser mesh than flat areas.

To detect these areas in a .STL file the density of the mesh can be estimated. Similarly to FEA, more complex areas will be represented by a denser mesh than flat areas. By placing voids near locations where the mesh is dense the likelihood of placing them somewhere that will cause a critical failure increases significantly.The ability to place a void completely inside of the part is a key part of a void attack, both for the functionality of the void and for avoiding detection. If a void is placed completely outside of the existing shell of the part it will serve no function and is not a concern. Similarly, a void that extends partially outside of a part will cause errors that will draw attention. The only dangerous void attacks are those where the void is completely enclosed. The task of placing a void inside of an arbitrary .STL file poses several challenges. First, .STL files may contain multiple shells.

The task of placing a void inside of an arbitrary .STL file poses several challenges. First, .STL files may contain multiple shells. A void may be fully enclosed inside of the part, but still intersect multiple shells causing an error. Second, because an .STL file may contain almost any geometry, they can be convex or concave.Because of this techniques for determining collisions based on convex geometries cannot be used. Other approaches, such as determining the center of mass, meet with similar difficulties as the center of mass may be outside of the object. In analyzing this problem a two-step method was found that could be used by an attacker to ensure that a void is placed completely inside of a part.Void ScalingAnother feature that makes a void attack more dangerous is the ability to scale the size of the void to better fit the part. The user has little concern from the placement of a void that will not have any harmful effects.

The user has little concern from the placement of a void that will not have any harmful effects. An attacker can increase the chance of failure by making the void larger, while still keeping it small enough to avoid detection. An attacker can easily modify the size of a void, however this increases the risk of it extending outside of the shell it is enclosed by. By combining scaling techniques with the previously discovered method for ensuring the void is completely enclosed, the chance that the void placed by the software would cause a failure was increased.The final attack software analyzes an .STL file, places a tetrahedron shaped void at either a random or semi-intelligent location, scales the void to a more idea size, and attempts to ensure that the void is fully enclosed inside of a single shell. Adding the void adds a total of four new facets to the file and creates a new negative volume shell consisting of those facets into an existing shell.

Adding the void adds a total of four new facets to the file and creates a new negative volume shell consisting of those facets into an existing shell. For a binary .STL file the total change in file size is 200 bytes since each facet requires 50 bytes to represent. Though this might be a discernable change in a small, simple part, it will be nearly invisible in a larger more complex part that has a file size of several megabytes.4. Effects of .STL Void AttackTo ascertain the potential impact of this specific attack, two experiments were performed. First the authors evaluated the effect of a “printed void” on the mechanical strength of a printed specimen.4.1 Effect of Voids on Part StrengthThe goal of the attack was to evaluate the impact of a void placed inside of a part for measurable degradation of part quality. To evaluate the effectiveness of a void at causing a failure an ASTM Standard D638-10 tensile test specimen, or dogbone, was used as a sample part.

To evaluate the effectiveness of a void at causing a failure an ASTM Standard D638-10 tensile test specimen, or dogbone, was used as a sample part. This part was chosen because it provided a straightforward way of measuring the quality of a part with and without a void.Figure 4 shows a cross section slice of a part with a void placed inside of it using the software. To quantify the effects of the void placement, finite element analysis was run to determine if the void was likely to cause a failure. Unsurprisingly, the Von Mises stress of the part is highest at the location of the void (Figure 5).In order to verify this result for parts manufactured using an AM process, several dogbones with and without voids were printed on via Powder Bed Fusion (a Sinterstation 2500 Plus machine) using Nylon 12 powder. Upon testing, all of the dogbones containing voids fractured at the void location, while the dogbones without voids failed normally, as shown in Figure 6.

Upon testing, all of the dogbones containing voids fractured at the void location, while the dogbones without voids failed normally, as shown in Figure 6. The average reduction in yield load was 14%, from 1085N to 930N, and the strain at failure was reduced from 10.4% to 5.8% as shown in Figure 7.Figure 4. Cross sectional slice of a dogbone infected with a void.Figure 5. Von Mises Stress of a dogbone infected with a void.Figure 6. On Left: Uninfected dogbones breaking at the neck. On Right: Infected dogbones breaking at the void location.12001000800)N( d a o L6004002000 0.00%2.00%4.00%6.00%8.00%10.00%12.00%StrainNo Void 1 Void1No Void 2 Void 2No Void 3 Void 3Figure 7. Load and strain data of parts with and without voids.In addition to the reduction in mechanical strength, the dogbones containing voids were tested to see if the voids were detectable through common inspection techniques.

Physical dimensions of the part revealed no differences, since the void was full enclosed. Weighing of the part did not reveal any noticeable differences since the void was small and filled with loose nylon powder, which had a similar density to the finished model material. The only effectiveway of detecting the presence of a void was to hold the dogbone up to a light source, revealing a darker area where the void was located; of course this technique is only useful for this thin (3.2 mm) part.4.2 RecommendationsThe results of this study show that more work is needed to protect AM systems against cyber-physical attacks. One of the great benefits of the .STL or .AMF file format is the ease in which files and designs can be freely shared. This is a great benefit to AM and would be lost if an attempt was made to move to a more closed “secure” format.

This is a great benefit to AM and would be lost if an attempt was made to move to a more closed “secure” format. Trying to create a “secure” replacement for the .STL would not only harm the field of AM as a whole, it would also fail to address the root of the problem. Any security system can be broken eventually, and replacing the .STL file would do little more than provide a false sense of security. A better way of addressing the threat of cyber-physical attacks is to develop better monitoring systems and procedures and to better train the workforce to detect these attacks.Improved Software Checks – Detection of shells with negative volumes, detection of very small shells, and detection of small features.Hashing/Secure Signing – Allow users to check that the file they receive is the same asthe one that was sent. Improved Process Monitoring – Monitoring process parameters through indirect “side channel” methods.

Improved Process Monitoring – Monitoring process parameters through indirect “side channel” methods. By observing physical parameters such as the melt pool temperature indirectly from machine settings such as laser power, attacks are more likely to be discovered.Operator Training – Better education of the workforce on the threats of cyber-physical attacks and on how to detect when an attack may be occurring.The first step that can be taken to prevent cyber-attacks is to improve the default softwarechecks that are run in common software programs used for checking .STL files. Existing software, such as Netfabb Pro, is easily capable of detecting shells with negative volumes.

Existing software, such as Netfabb Pro, is easily capable of detecting shells with negative volumes. However, its default error check does not present the user with an alert when a negative volume shell is detected, even though in most parts negative volume shells represent an error or attack on the file.A second check that can be performed is a test for shells with very small volumes since theyare unlikely to be important features, and are often unneeded artifacts left over from another process like 3D scanning. Testing for these shells increases the chance of detecting a void based cyber-attack in addition to alerting the user of potentially unwanted or unneeded data.Finally, more research is needed into detecting small features that are not independent shells. Very small features that cannot be built by an AM process may indicate the presence of a cyber- attack. Even if the features are not malicious, the user should be alerted that they will not be printed correctly.

Even if the features are not malicious, the user should be alerted that they will not be printed correctly. The inclusion of these three checks would not only make detecting an attack easier, but would also help the user in identifying potential problems in their models.Hashing is a technique commonly used in security to ensure the validity of a file. The file isrun into the hashing function, which generates a string of character called a hash. The hash is then posted along with the file. When a user downloads a file they can run it through the same hashing function and compare the resulting hash with the posted hash. If the two hashes match,the file can be assumed to be identical to the original. What makes hashes effective is their ability to convert a large file into a simple string that can be easily shared. Any small change in the file generates a large change in the hash.

What makes hashes effective is their ability to convert a large file into a simple string that can be easily shared. Any small change in the file generates a large change in the hash. The simple adding of a single character to a text file will completely change the hash that is generated.At the most basic level hashes can be (and are) used to ensure that the .STL file received hasnot been tampered with. While this does increase security it adds some additional work to the process. Additionally, a file could potentially be attacked after it has been received and hashed, or before it was hashed to begin with. This problem could be addressed in part by including the hash function at the time of file creation, within the CAD software itself, and generating a hash function at the last step in the process, where the .STL file is loaded in to the printer software to be converted to a toolpath.

One area of further research is the creation of a physical hash, a hash that incorporates elements from both the physical and the cyber side to give added resilience against cyber-attacks.Improved process monitoring is an ongoing goal and area of research for additivemanufacturing in an effort to better control and improve them. Some solutions have already been explored, such as using optical sensors to provide closed loop control for layerwise laser melting (LLM) [10]. Indirect measuring, or “side channel” measurements, such as measuring the temperature of the melt pool to determine the laser power instead of simply asking the machine can be effective at detecting cyber-attacks effecting the machine parameters, since, as Stuxnet demonstrated, a clever attack can cause a system to report false data.While feedback systems provide valuable information for process control, used alone they are insufficient to detect cyber-attacks.

If the .STL file for a part is infected with a void a closed loop system will notice the void, compare it to the part file, and conclude that it is an intended design feature. For this reason more research needs to be done to establish statistical process control (SPC) for AM systems that works to detect the effects of cyber-attacks. By combining side channel measurements and using SPC to establish baseline operating parameters, systems can be made more robust and also more resistant to cyber-attacks.The final area that needs improvement is in education and training. The vast majority of malicious software is installed or transmitted unwittingly by uneducated users. Educated workers are more likely to detect and prevent a cyber-attack. Informing students and workers about the potential risks of cyber-attacks can prevent these attacks from occurring and help diagnose them faster when they do occur.5.

Informing students and workers about the potential risks of cyber-attacks can prevent these attacks from occurring and help diagnose them faster when they do occur.5. ConclusionsWith the increasing number of manufacturing systems connected to networks, more workneeds to be done to ensure the safety of these systems. Additive manufacturing systems in particular have unique vulnerabilities presented by the ability to affect the internal layers without affecting the exterior. An overview of the AM process chain (Section 2) showed that the .STL file was the most vulnerable attack vector due to its universality and ease of editing. Further investigation into the .STL file revealed a method by which a void could be placed inside a part, while avoiding detection by common process checks (Section 3). This void placement was demonstrated to cause a 14% reduction in yield load in a tensile test specimen.

This void placement was demonstrated to cause a 14% reduction in yield load in a tensile test specimen. Based on the results of this study, it appears that a real threat from cyber-attacks exists and that further research needs to be done on how to mitigate such attacks. The inclusion of software checks,hashing, process monitoring, and worker training are proposed as methods of reducing these threats. Future work includes the development of physical hashing techniques and of improved side channel process monitoring and controlReferences[1] Evans D. The Internet of Things: How the Next Evolution of the Internet is ChangingEverything. CISCO White paper; 2011. Available from: https://www.cisco.com/web/about/ac79/docs/innov/Io-T_IBSG_0411FINAL.pdf [Retrieved March 13, 2014]..[2] Bayuk J, Cavit D, Guerrino E, Mahony J, McDowell B, Nelson W, Snevel R. Malware Risksand Mitigation Report. Washington, DC: BITS Financial Services Roundtable; 2011. [3] Watin-Augouard M.

CLIC OLEFIN COPOLYMER STRIP PROCESSING FOR FREEFORM FABRICATION OF MULTI-LAYERED MICROFLUIDIC SENSING SYSTEMSAymen BenAzouz*†, Richard O'Connor†, Mercedes Vasquez*, Dermot Brabazon*†, and Brett Paull* *Irish Separation Science Cluster, National Centre for Sensor Research, Dublin City University †School of Mechanical and Manufacturing Engineering, Dublin City University, IrelandAbstract Researchers have developed techniques for multi-layered fabrication of microfluidic chips which allow for increased scope of channel geometries and associated improved sensing capabilities. In these techniques, slits have been fabricated in thin layers of polymer or glass, typically of a couple of hundred micrometers thick. These layers are then bonded to each other using adhesives, hot embossing or a combination of bonding methods.

These layers are then bonded to each other using adhesives, hot embossing or a combination of bonding methods. This paper presents a new fast freeform methodology for 3D channel geometries to be fabricated in COC chips using laser and xurographic processing for slit formation and cyclohexane promoted bonding for multi-layer joining at room temperature.IntroductionMicrofluidics devices are widely used in research and in chemical, environmental and biotechnology sensing applications. A microfluidic device or chip can be a device of a few millimetres to a few square centimetres in size, integrating 3D micro and nano features. They can vary from single channel per chip layouts to much more complicated designs. They are fabricated to handle small quantities of fluids, have high sensitivity, and their function may incorporate several laboratory functions in a single chip [1].

They are fabricated to handle small quantities of fluids, have high sensitivity, and their function may incorporate several laboratory functions in a single chip [1]. Their development has introduced a positive transformation within chemical and biological research due to their rapid and high-throughput analysis and minimized consumption of sample and reagent [2]. Traditional methods of chip fabrication involve bonding two polymer or glass layers after engraving the desired design on one of the layer surfaces. Lithography, electron beam writing, photo-patterning and chemical etching are the most common engraving techniques used to fabricate microfluidic chips [3]. Thermal bonding is the most widely used and established method of bonding.

Thermal bonding is the most widely used and established method of bonding. Other bonding techniques include chemical treatment of surface, application of adhesive layer and laser exposure to the active bonding medium [4].Fabrication and utilisation of multilayer microfluidic chips is an emerging area of work and has become a focus of the microfluidic research field in recent years [5]. This technique comprises the process of slit fabrication in thin layers of glass or polymers, typically of a couple of hundred micrometers maximum thick and then the bonding of these layers to form a 3D structure inside the microfluidic chip, see Figure 1.Figure 1 Schematic of the process of multi-layer microfluidic chip fabrication.At the beginning of last decade the first attempts to fabricate multilayer chips, were made with Polydimethylsiloxane (PDMS), silicon, glass or a combination of two or more of these materials [6-11].

Since most of these works are focused in the fabrication of microfluidic devices for specific chemical or biological application, testing was limited to flow and chemical capabilities aspects. In more recent publications, up to eight-layer microfluidic chips of poly(methylmethacrylate) (PMMA) have been fabricated [5, 12]. Pressure tests, fluidic electrical resistance and mixing experiments, have been performed to assess the strength of the bonding as well as the electrical and chemical capabilities of the chips. All these publications used photolithography, wet-chemical and ion etching techniques to fabricate slits, except in some works where CO2 laser was used to create features in PMMA [12].Despite limited research with solid freeform techniques (SFF), it is easy to see that there is a high potential to develop 3D microfluidic devices using 3D SFF microfabrication methods. Bonding remains the most critical and inconsistent step in multilayer microfluidic chip fabrication [2].

Bonding remains the most critical and inconsistent step in multilayer microfluidic chip fabrication [2]. SFF methods can overcome this problem; methods such as 3D printing circumvent traditional bonding techniques leading to microfluidic devices with high integrity and precise alignment of feature. In this work, a new fast freeform process of fabrication of multilayer Cyclic Olefin Copolymer (COC) microfluidic chip is presented.Materials and methodsThe COC was provided by Zeon Chemicals Europe Ltd, in the form of A4 sheets. COC compositions used in this work were ZF14 and ZF16. Layer thickness of 40µm, 100 µm and 188µm were processed into 3D structures.

COC compositions used in this work were ZF14 and ZF16. Layer thickness of 40µm, 100 µm and 188µm were processed into 3D structures. These materials were chosen as they are becoming increasingly popular due to their advantageous properties of high optical transmission, low refractive index, low water absorption and resistance against eluting chemicals such as methanol, ethanol, acetone and nitric acid [13].Figure 2 presents the straight channel test design fabricated in this work. The chips were assembled from three layers.

The chips were assembled from three layers. The chip dimensions were 40×20mm, the inlet and outlet ports were 1mm in diameter, while the channel length was 30mm.Figure 2 Schematic of the microfluidic chip single channel test design.The main steps of fabrication of the microfluidic device in this work were the following:creating slits on the thin layers of COC • exposing layers to Cyclohexane vapour • aligning slits and applying pressure for determined timeThe characterization and testing consisted in flow tests to analyse the applicability of the chips and tensile tests to measure the strength of the bonding.Cyclohexane bondingA bonding technique previous published was used in this work [2]. The COC layers were bonded by exposing them to cyclohexane vapour. The apparatus used included a hotplate, a number of clock glasses, an alignment jig and a pneumatic press. Before bonding, the COC was cleaned and rinsed with isopropanol and water, then dried with nitrogen.

Before bonding, the COC was cleaned and rinsed with isopropanol and water, then dried with nitrogen. The protective polyethylene (PE) cover layer was removed using scalpel and tweezers, the sample were rinsed and dried once more before exposure to cyclohexane. Using the hotplate, the cyclohexane was heated to 70ºC to produce vapours within the chamber. Once there was a uniform vapour presence within the chamber the COC was exposed to the solvent by swapping the chamber cover with another that had the COC sample attached to its underside.After exposure to the solvent the samples were placed in a pre-made alignment jig. Half of the surface area was covered with card to prevent bonding of the entire surface area. After pressing, the newly bonded samples were then exposed to UV light (2 J/cm2, 260 nm) to increase bond strength. Nine sets of processing conditions for bonding were examined.

After pressing, the newly bonded samples were then exposed to UV light (2 J/cm2, 260 nm) to increase bond strength. Nine sets of processing conditions for bonding were examined. These include three periods of vapour exposure (30, 40, and 50 seconds) and three periods for applied bonding pressure of 0.4 MPa (3, 4, and 5 minutes), see Figure 3. These nine treatment conditions were tested for bond strength, see Figure 4. The COC samples prepared for tensile testing were ZF14, 40×20mm and 40 μm thick.(a)(b)Figure 3 Pictures of (a) sample being exposed in clock glass and (b) compression in the pneumatic rig at 4 bar.Tensile testingTensile testing of the bonded samples was performed using a Hounsfield H20K-W tensile testing machine. The rate of extension was maintained at 2.5mm/min. The samples were placed in the jaws of the machine as shown in Figure 4.

The rate of extension was maintained at 2.5mm/min. The samples were placed in the jaws of the machine as shown in Figure 4. The initial distance between the jaws was set at 20mm for each test; the force versus extension data were recorded for each test. These were then used to calculate the stress and percentage strain experienced at the bond surface.(a)(b)Figure 4 Picture of the (a) start of a tensile test of the bond strength and (b) a sample tested to failure.Microfluidic chip fabricationThe same procedure for tensile test sample bonding was used to bond the test microfluidic chips. To reduce the risk of damaging the more delicate middle layer, the top and bottom layers were simultaneously exposed to the cyclohexane vapour. After exposure the layers were assembled using the same jig as previously stated. Once aligned the chip was pressed by hand to push any possible air pockets out of the bonded areas and then placed in the pneumatic press.

Once aligned the chip was pressed by hand to push any possible air pockets out of the bonded areas and then placed in the pneumatic press. After pressing, the newly bonded samples were then exposed to UV light.Slit fabricationThe feature on the chip layers were produced by two methods, CO2 laser ablation and razor writing. These methods were chosen as they require little initialization, with easy reproducibility and the can produce multiple layers within one set up. The chips were assembled from three layers, as shown in Figure 2. The chip dimensions were designed as 40×20mm, the inlet and outlet ports were 1mm in diameter, while the channel length was 30mm.Laser processingA 10.2µm wavelength continuous CO2 laser, GEM-100L, was used with the chip layout imported from an AutoCAD design. The laser processing was optimised to reduce excessive burning of the material. It was found that a cutting speed of 2mm/s and a power output of 7 W produced the most defined channels.

The laser processing was optimised to reduce excessive burning of the material. It was found that a cutting speed of 2mm/s and a power output of 7 W produced the most defined channels. A constant flow of compressed air was maintained on the cutting surface during ablation which improved the channels smoothness by cooling the molten polymer quickly. The channel widths produced using this method were on average 180µm. A polyethylene cover on the COC film was removed after laser ablation to reduce contamination and scratching of the material. It was found that placing the COC film with the PE cover facing up during laser cutting made removing the cover from the cut layers much easier, cutting with the PE facing downwards made the removal of the cover difficult and also lead to fibres of PE becoming trapped in the burring from cutting. The time taken to produce nine layers, which corresponds to three assembled chips, was 15 minutes.

The time taken to produce nine layers, which corresponds to three assembled chips, was 15 minutes. The microfluidic chips produced by laser ablation were made from ZF14 with film thickness of 100µm.Razor writingThe razor writing or xurographic method was much less time consuming. The Graphtec Craft ROBO Pro, which has a specified mechanical positioning resolution of 5µm, was used to cut the individual layers for the multilayer chips. It was possible to populate a much larger area with chip layers as the plotter was capable of carrying film sizes of up to A3, when compared to the limited 150 by 115 mm area possible in the laser ablation machine. The chip patterns were again imported from AutoCAD, this was the same as used for laser ablation. An A4 sheet of the Zeonor ZF16 with a thickness of 188µm was cut into 30 layers (providing material for 10 assembled chips); the total cutting time required was approximately 5 minutes.

An A4 sheet of the Zeonor ZF16 with a thickness of 188µm was cut into 30 layers (providing material for 10 assembled chips); the total cutting time required was approximately 5 minutes. The A4 sheet of COC was first attached to an A3 adhesive carrier sheet, it is important to place the COC with its PE protective cover facing downwards as this prevents PE fibres and debris from attaching to the blade, which would reduce the accuracy of the cut. With the rotary style of cutting relatively little heat was introduced during the cutting avoiding the burring potential present in laser ablation. The depth of the cutter can be varied if thicker material would need to be cut, however different cutters are also required for the cutting of certain materials. The blade used in this work (CB09UA) was suitable for cutting material of thicknesses up to 250µm. The cutting speed at which the plotter operated was 20mm/s.Flow testingThe microfluidic chips fabricated were subjected to flow tests.

The cutting speed at which the plotter operated was 20mm/s.Flow testingThe microfluidic chips fabricated were subjected to flow tests. Droplets of red-diluted water were deposited on one of the inlets of the chip and vacuum pressure applied on the outlet. Zeonor ZF16, required a longer period of exposure to cyclohexane vapour (60 and 75 s compared to 30 and 40 for ZF14; 50 s was not used for ZF14 in these tests). This can be explained by the different glass transition temperature of the two materials. The duration of pneumatic exposure was maintained at 5 min for all samples. Additional bonding strength through 2W UV exposure was applied to some samples.Results and DiscussionBonding strength testing Figure 5, 6, and 7 show the tensile test results for the cyclohexane exposure times of 30, 40 and 50 seconds respectively.

The increased exposure time allowed greater penetration of the cyclohexane vapour into the bulk material, allowing improved activation and cross linking of polymer chains. There is however a disadvantage to overly long exposure time as the material softens, increasing the likelihood of blocking channels during bonding. Saturation of the COC with cyclohexane will cause rapid softening and to some extent degradation of the material. It was noticed that a transfer of some COC material occurred on equipment that was used to handle the material when the exposure time exceeded 50 seconds. A confidence interval was constructed to compare the UTS achieved from different solvent exposure durations. At 90% confidence, the UTS for the chips bonded at 30 seconds solvent exposure were significantly less strong than chips bonded at 40 and 50 second exposure times.

At 90% confidence, the UTS for the chips bonded at 30 seconds solvent exposure were significantly less strong than chips bonded at 40 and 50 second exposure times. No significant difference was noted between bond strengths for chips bonded with 40 and 50 s exposure times.Tearing of a couple of the COC films occurred during tensile testing. As the bonding in areas around the edges of the samples was consistently poorer than internal bonding, this could cause these couple of thinner material samples to fail at the interface between bonded and non- bonded regions. The bonding process, including UV cross linking, likely strengthens the material, while the compression causes it to become denser. The alignment of the samples during bonding, as well as during testing may also explain the tearing phenomenon experienced.

The alignment of the samples during bonding, as well as during testing may also explain the tearing phenomenon experienced. Incorrectly aligned sample, when under tension in the manner in which it was performed, can produce a small torque/twist of the material, causing the tensile force to concentrate at a corner of the bonded area instead of it being perpendicular to it.By comparing the stress strain charts it can be seen that an increased exposure time results in an increase in the bonds tensile strength, however the same cannot be said for increased compression duration. The stress strain curves also demonstrate the variability of the bonding strength across the length of the bond area. The best example of this is seen in Figure 7 where the samples have been exposed for 50 seconds. The red and yellow curves, which relate to a 3 and 5 minute compression duration, are not smooth in nature. The stress increased and decreased as strain was increased.

The red and yellow curves, which relate to a 3 and 5 minute compression duration, are not smooth in nature. The stress increased and decreased as strain was increased. This would most likely have been caused by the variation in bond quality, and the presence of voids caused by air pockets and foreign particles.Slit formationThe repeatability of the laser ablation and blade cutting methods of rapid prototyping are well documented [14]. In this application the surface finish of the material is critical to a device’s success. Edge burrs which can hamper proper bonding, can be caused on the underside of the film during laser ablation. For the fabrication of multilayer chips from polymer films the blade cutter is a more economical and time effective method, while eliminating the burring caused by laser ablation. The plotter itself is no larger than a typical desktop ink jet printer, with a similar user interface. However it is limited to polymer films of thickness of 250µm or less.

The plotter itself is no larger than a typical desktop ink jet printer, with a similar user interface. However it is limited to polymer films of thickness of 250µm or less. Laser ablation has a greater flexibility in the materials (such as glass) and thickness that can be used for chip fabrication. Some ablation was performed on glass using the CO2 laser, with varying degrees of success. In order to perforate the glass the lasers power must be of sufficient intense, however rapid increase in temperature can lead to cracking of the material by thermal shock.(a)(b)(c)(d)Figure 6 Pictures of (a) 1 mm diameter laser cut inlet, (b) 180µm width laser cut channel, (c) 0.7 mm diameter blade cut hole, and (d) 50 µm blade cut channel.Flow testingQuality of the bond was tested during the fluid flow tests. Chips with more voids or foreign material were noticeably less leak resistant.

Chips with more voids or foreign material were noticeably less leak resistant. This could be attributed to the burring and the excess material around the edges of each layer as produced by laser ablation. This however does not explain the poor bonding of the chips produced using the blade plotter. The most likely explanation for this is the time taken in alignment and the method of alignment itself. After exposure to the cyclohexane vapour the bottom layer of COC was removed from the clock glass, placed in the alignment jig, the middle layer was placed on top followed by the top layer. In order to complete the chip the inlet and outlet holes must be correctly aligned with the channel, all three layers must then be removed and pressed by hand to remove air pockets trapped in the bonded area.

In the time taken to align all three layers and remove them from the press the cyclohexane is continuously diffusing further into the material and evaporating from the surface The process itself also introduced unwanted COC particles into the bonding area, while the bonding surfaces can be damaged by over stressing the tacky COC, creating large particles which in turn cause large air pockets.After bonding, the microfluidic chips were tested to investigate flow through the channels and to assess leakage between the layers. Figure 9 shows a representative selection of the chips produced. Figures 9 (a) and (b) show the same chip before and after a coloured dye was applied. The inlet and outlet holes and the channel can be seen in Figure 9 (a), as well as the bonded areas. The non-complete sealing of the channel is highlighted in Figure 9 (b) after dye was applied to one of the holes and leakage occurred immediately without the need to apply a vacuum.

The non-complete sealing of the channel is highlighted in Figure 9 (b) after dye was applied to one of the holes and leakage occurred immediately without the need to apply a vacuum. Some flow was directed through the channel after the vacuum was applied. Similarly, Figure 9 (c) shows poor bonding surrounding the channel, this was as a result of the burring caused by laser processing. When a vacuum was applied to this chip little or no flow was seen within the channel, it instead flowed around it. There is better overall bonding displayed in the chip in Figure 9 (d). This shows the most successful attempt at multilayer chip fabrication and assembly. As can be seen the dye filled the entire channel. This was accomplished with the need for a vacuum pump, due to capillary attraction. After a period of time the solution began to leak into the area surrounding the channel.

This was accomplished with the need for a vacuum pump, due to capillary attraction. After a period of time the solution began to leak into the area surrounding the channel. This does however demonstrate that multilayer microfluidic devices can indeed be fabricated using COC films and cyclohexane vapour assisted bonding. This last channel was fabricated at a much higher compression pressure.ConclusionsFrom this investigation it was found that the optimum duration for exposure of COC to cyclohexane vapour was 40 seconds. This gave a good bond UTS but also maintains more of the materials integrity during bonding, reducing the chance of blocking the channel as compared to 50 second exposure time. Alignment of the layers is of utmost importance to the performance of fabricated microfluidic chips. To achieve this, automating the procedure should be considered which would reduce over stressing/damaging of the material and the infiltration of the bond area with foreign objects.

To achieve this, automating the procedure should be considered which would reduce over stressing/damaging of the material and the infiltration of the bond area with foreign objects. With improve alignment, and reduced handling time, comes improved bond quality. By reducing the time taken to align the layers and place them for pressing, more cyclohexane will still be present on the materials surface, increasing the effectiveness of the process.

By reducing the time taken to align the layers and place them for pressing, more cyclohexane will still be present on the materials surface, increasing the effectiveness of the process. To continue to improve the performance of the COC multilayer chip the blade cutter and laser process need to be used to prepare the individual layers as they have great potential to be even faster and provide more effective methods of multi-layer freeform fabrication.(a)(b)(c)(d)Figure 7 Pictures of microfluidic chip sample (a) before flow test, (b and c) after flow test with leaking and (d) after successful flow test.AcknowledgementsThe authors would like to thank Science Foundation Ireland (Grant Number 08/SRC/B1412) for research funding under the Strategic Research Cluster programme.References1.2.3.4.

data format providing side wall orientation and adaptive slicing for use in stepless rapid prototypingWei Fang, Chan Siang Ngooi, and Haiqing Gong* School of Mechanical and Production Engineering Nanyang Technological University Nanyang Avenue Singapore 639798 Email: mhqgong@ntu.edu.sgAbstractCurrent Rapid Prototyping systems mainly use 2D layers building techniques that result in the 'staircase' effect on slanted surface. A new method has been developed to eliminate the 'staircase' effect and to improve the surface quality by extending the 2D layers to 3D layers building. In this approach, a new data format, Layer Transfer Interface (LTI), is introduced to It provides a faster slicing algorithm and accurate generate layers having slanted side wall. reconstruction of 3D objects. This format is independent of any particular RP machines.

reconstruction of 3D objects. This format is independent of any particular RP machines. Furthermore, adaptive slicing has been achieved using this format and implemented on a five axis milling RP system.1.IntroductionResearch in rapid prototyping places a constant emphasis on the need for improvement of accuracy and surface smoothness. From the CAD model point of view, the three main ways of slicing are:Fixed slicing where peaks or flat areas (areas parallel to the slicing planes) are ignored [1,2,3,4].Adaptive slicing where the slice thickness varies in response to the surface curvature [2,3,4].Direct slicing from CAD original models as opposed to tessellated models [5,6].Presently, Common Layer Interface (CLI) [7] and 3D system's format SLC [8] are the universal slice formats that are simple, efficient and unambiguous. However, the staircase effect cannot be totally eliminated using these formats.

However, the staircase effect cannot be totally eliminated using these formats. Also, there are redundant data between slices due to 2D slice contours in these formats, which result large data file being produced.The use of slanted and ruled slices can remove the staircase effect. Slanted slices can be generated using these approaches [9]. However, it requires a minimum of 4-axis control to make the slanted edges of the RP part. In order to have a slanted slice, a data format must be able to be read by the RP machines... To whom the correspondence should be addressed6212.Layer Transfer Interface (LTI) FormatA new data format is proposed to use in new RP machines with at least 4-axis control. Figure 1 shows the slope information of a point of a slice.

Figure 1 shows the slope information of a point of a slice. These additional information of the roll and pitch angles enable slanted slices to be represented.xyz is the local axis frameZ Upper SliceLower SliceFigure 1 - Slope information of a pointThe LTI data format consists of a collection of segments as shown below.II LTI format Parts_BeginHeadNumber_of_Layer: [number] Part_size: [xl, yl, zl, x2, y2, z2] units: [mm or inch]End Head Layer_BeginIndex: [number] Thickness: [z] Number_oCContour: [number] Contour_Beign:Number_oCHole: [number] Begin_Tool:Number_of_Point: [number]X, y, z, <X, p,End_Tool Hole_Contour_Begin:Begin_Tool:Number_of_Point: [number]x, y, z, <X, p,End_ToolHole_Contour_EndContour_End622The first segment is the header segment that contains the global information of solid model such as layer number, size etc. The second segment contains the layer contour data.

The second segment contains the layer contour data. Inside this segment, the layer structure is sectioned into many parts that consists of contour, hole and hatching data for each slice.With the slope information of the LTI, data transfer to other slice data format can be done readily. In the generation of the sliced contour data for a user defmed layer thickness, the interpolation from the upper and lower contours can be carried out. This technique is fast and easy and also adaptive slicing can be implemented.3. MethodA new slicing method is developed to provide the roll and pitch angle information. Currently, only faceted model is used. The slicing starts with the determination of the topological information of the input model. This topological information consists of the vertex and face lists. The vertex and face lists are linked using a circular link list. The vertex list is later sorted along the direction ofbuild direction.

The vertex and face lists are linked using a circular link list. The vertex list is later sorted along the direction ofbuild direction. Using the vertex information, the flat areas are determined.Next, an advanced marching algorithm is carried out using two parallel planes to obtain the slope information. With this information, the Layer Transfer Interface (LTI) data file is then generated. This data format is independent of any RP machines. A five-axis milling machine has been developed to use this data format to build parts using layering techniques. This data format is further preprocessed to create each layer with variable layer thickness. The system parameters that affect the slice thickness are the tool diameter and tool length. Figure 2 shows the faceted .STL model. The slicing algorithm is developed using C++ with OpenGL working in Windows NT platform.

Figure 2 shows the faceted .STL model. The slicing algorithm is developed using C++ with OpenGL working in Windows NT platform. Figure 3 shows the sliced part with slope information for slanted side wall using LTI data format.Figure 2 - A .STL facet model623Figure 3 - A sliced part with slope information for slanted side wall using LTI format4.ImplementationA five-axis milling machine was developed to build layers with slanted side wall. Figure 4 shows a part built using this machine. The slope information obtained from the LTI format was used to machine the slanted side wall. With this slanted building technique, part accuracy and speed are greatly improved. The offset and tool path calculation are performed using this new LTI format.Figure 4 - A part with layer having slanted side wall fabricated using a five-axis milling machine6245.ConclusionThe new data format (LTI) was proposed for producing stepless RP parts.

This format increases the part accuracy and it is easy to perform adaptive slicing. With the capability of the RP machines keeps on improving, the use of more than 3-axis control to built layers will be seen in the near future.6.References[1] Wozny, M.l 1992. Systems issues in solid freeform fabrication. Technical Report 92023,Rensselaer Polytechnic Institute, Troy, NY12180.[2] Sub, Y., and Wozny, M.l 1994. Adaptive slicing of solid freeform fabrication processes.Solid Freeform Fabrication Symposium Proceedings. University of Texas at Austin, Austin, Texas. pp. 404-411.[3] Dolenc, A., and Makela, I.. 1994. Slicing procedure for layered manufacturing techniques.Computer-Aided Design. Vol. 26:2, pp. 119-126.[4] Kulkarni, P., and Dutta, D. 1995. Adaptive slicing for parametrizable surfaces for layered manufacturing. Design Automation Conference Proceedings. Boston, Massachusetts. pp. 211-217.[5] Guduri, S., Crawford, R.H., and Beaman, J.1. 1992.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA Data Integration Framework for Additive Manufacturing Big Data ManagementMilica Perišić1, Dimitrije Milenković2, Yan Lu3, Albert Jones3, Nenad Ivezić3, Boonserm Kulvatunyou31Faculty of Organizational Sciences, University of Belgrade, Serbia 2TX Services, Serbia 3Systems Integration Division, National Institute of Standards and Technology, Gaithersburg, MDAbstract Large amounts of data are generated throughout the entire, AM, part-development lifecycle. Data are generated by various functions within process monitoring, material characterization, equipment status, and part qualification. Hence, data integration and management are critical in streamlining, accelerating, certifying, and deploying these functions.

Hence, data integration and management are critical in streamlining, accelerating, certifying, and deploying these functions. However, achieving that integration and management has several challenges because AM data embodies the four characteristics of Big Data - volume, velocity, variety, and veracity. This paper proposes an AM framework as a foundation for addressing those challenges. In the framework, AM data are streamed, curated, and configured automatically for real-time analysis and batch processing, which increases the effectiveness of archiving and querying that data. The framework also includes a description of the associated AM metadata, which links the various data types and improves browsing, discovering, and analyzing that data. Finally, the framework can be used to derive requirements for standards that enable data sharing.1. Large amounts of data are generated through the entire AM development lifecycle.

Finally, the framework can be used to derive requirements for standards that enable data sharing.1. Large amounts of data are generated through the entire AM development lifecycle. Data is generated and collected for material characterization, process monitoring, part qualification, etc. Hence, integrating and managing this data is critical in streamlining, accelerating, certifying, and deploying those functions. However, successfully integrating and managing this data introduces several new challenges because AM data embodies the four characteristics of Big Data - volume, velocity, variety, and veracity. Moreover, addressing these challenges is critical for advancing the capabilities and use of AM production technologies. Today, AM technologies are mainly used for rapid prototyping; but they are slowly emerging as a commonly used production technology in several industry sectors [1]. AM data integration is one of the top roadblocks to completely automated production management.

AM data integration is one of the top roadblocks to completely automated production management. A major reason for this roadblock is the lack of established methods and standards that allow a quick, “plug-N-play” type of integration of AM data sources with various manufacturing and enterprise applications. For example, the large amounts of high-speed, in-process, monitoring data such as melt pool images cannot be acquired, and automatically processed for real-time, or near-real-time control.IntroductionA “divide and conquer” approach is proposed in this paper to create a data integration framework that addresses these issues. The framework has seven functional steps, which are discussed in the next section, including Defining Dataset/Data Source, Collecting Data, Queueing Data, Archiving Data, Downgrading Data Amount, Building Decision Models, and Using Decision Models.

In addition, the paper offers a guide for using the framework with the current industry standards.14192.A Big Data Integration FrameworkFigure 1 presents a framework that can facilitate the creation of a big data integration system that would enable real-time monitoring and control and long-term data archiving for offline analyses. The framework consists of seven steps. Step one provides a clear definition of data that comes to the system and ensures that future misunderstandings between the data provider and the system maintainer would be avoided. Step two collects or receives the data from the data provider and enters it into the system to be further processed. Step three prevents the system from being overwhelmed by queueing raw data until it can be processed and stored in the system. Step four decides how the data is stored, which is the key to having a stable and tenable system that makes data accessible both using queries and bulk downloads.

Step four decides how the data is stored, which is the key to having a stable and tenable system that makes data accessible both using queries and bulk downloads. Step five achieves a sustainable system by reducing and managing the data size and saving only the data necessary for future use cases. Step six creates new value by building Decision-Making algorithms and models that use the stored data as inputs. Step seven creates programs that automatically allow use of the system through alarms, monitoring reports, and automatic actions. Each step is discussed separately with the aim to emphasize the new standards that are needed.1234567Figure 1. A big data integration framework2.1 Defining a Data Source A clear definition of a data source and its data is critically important both for the Data Provider and for the System Maintainer because it ensures that both parties have the same understanding of that data.

Such a definition represents a joint agreement between the Data Provider and the System1420Maintainer. Moreover, if this definition is not standardized and defined, it can cause misunderstandings throughout the whole life cycle. Three important descriptions are needed to create a standardized definition.2.1.1 Data-Source description: Data-source description provides the information about the device that generates the data as well as the measurement associated with the data. “Data Source Name” specifies the name of the data source. “Data Source Type” is necessary to indicate its physical characteristics. For example, the “Data Source Type” can be “Coaxial Camera-based MPM”, or “Layerwise Overview Imager” etc. “Device Type”, “Device ID”, “Device Manufacturer”, “Device Model” and “Device configuration” captured the metadata of the measurement devices or actuation devices. In the case of the co-axial-camera-based, melt-pool, monitoring system, the device type could be “CMOS Camera”.

In the case of the co-axial-camera-based, melt-pool, monitoring system, the device type could be “CMOS Camera”. The metadata definition for data sources can be leveraged on Dublin Core Metadata Initiative schema1.2.1.2 Data description: Another block of information is about the data itself. Typical attributes include “Data Category”, e.g., Sample, Event, or Condition based on MTConnect2. Data Type can be Value, TIME SERIES, and DATA SET as defined in MTConnect. Additional data types include images, videos, and 3D models, etc. Coordinate System is also a data field that could be very important to interpret position and size of measurement data.2.1.3 Load description: To integrate data sources, it is important to know whether the data are from streaming or batch upload as well as additional metadata about the load, e.g., size and frequency. For example, consider an image data type.

For example, consider an image data type. Will an image file and its metadata be sent as one object or two? What’s the expected size? Approximately, how many data instances will be sent per day or per batch? What’s the communication protocol? In the case of streaming data sources, some additional information is needed, such as sample interval and streaming triggering mechanism.Integration of streamed data is more challenging compared to the batch upload, so that will be the focus of the paper. Nevertheless, the framework can be applied to batch uploads as well.2.2 Collecting Data After defining the data, it is necessary to decide how the data will be collected from the data source. In most processes, including Additive Manufacturing processes, in-process data is created by following some of these three policies: sample, on event, and on condition.

In most processes, including Additive Manufacturing processes, in-process data is created by following some of these three policies: sample, on event, and on condition. This means that data will be created during a defined sample interval, when a defined event occurs, or when a defined condition is satisfied, respectively. Independently of how the data is created, there are two main approaches when it comes to the data collection – push or pull. These approaches can be applied for both streaming and batch data processing.In the first case, the Receiver Web Server, shown in Figure 1, is a data gateway. At this step, it is important to define a standard for exchanging the data. The Receiver Web Server is an application written in Java or some other programming language that receives the data through HTTP/HTTPS protocols and forwards it to the queue.

The Receiver Web Server is an application written in Java or some other programming language that receives the data through HTTP/HTTPS protocols and forwards it to the queue. Queue and its purpose are discussed in the next step of the framework.In the second case, when the data provider, the device, cannot push the data to the system, the approach consists of developing an application that will pull data from the data provider’s specific1 https://dublincore.org/schemas/ 2 https://www.mtconnect.org/standard201811421location and forward it to the queue. The best practice is to define a standard location for both, the data, such as images created during an AM process, and its metadata. The application can use some of the standard protocols to pull data such as SSH and FTP. The application can be created using Shell or any other scripting language and can be scheduled using CRON3.

The application can use some of the standard protocols to pull data such as SSH and FTP. The application can be created using Shell or any other scripting language and can be scheduled using CRON3. Also, for this case some recent technologies could be used, such as Apache NiFi4.2.3 Queueing Data Processing a data instance requires system resources. To prevent the system overload, a message queue can be helpful. In that way, data will be temporarily stored in the queue until it is processed and stored in the system. Multiple queues may also be used. For example, the first queue stores raw, unprocessed data waiting to be processed. The second temporarily holds processed data waiting to be stored in the system. Common message queue technologies used for these purposes are Apache Kafka5 and IBM MQ6.2.4 Archiving Data Deciding on which persistent-storage technology to use is key to having a stable and tenable system. Metadata and image data should be defined and stored separately.

Metadata and image data should be defined and stored separately. It is common to use file systems as image storage, but metadata can be stored in a database that can be searched and queried quickly and easily. Following that approach, metadata will have one additional field - image location from where the image can be read.2.4.1 Metadata Storage: Some of the options for storing metadata are a relational database, a file system (e.g., Hadoop Distributed File System - HDFS4), or a document database (e.g., MongoDB7). Relational database management systems (RDBMS) are the industry standard. The main advantage of RDBMS is that most engineers know this technology and there are integrations with many external tools. Considering that metadata will not change after insertion, document databases can be used as well. However, the file system is appropriate only if data queries are based on few fields with a limited number of values.

However, the file system is appropriate only if data queries are based on few fields with a limited number of values. When this is the case, these fields and values can be used to partition the metadata within a file system. This approach allows faster metadata access.2.4.2 Image Storage 2.4.2.1 File System Storage: One way to store big data, including images, is to use a file system. Currently, a commonly used one for big data is Hadoop Distributed File System (HDFS8). HDFS is highly fault-tolerant and is designed to be deployed on low-cost hardware. In recent years, many other tools and solutions were developed to support Hadoop core components. All of them together make data analysis, storage, and maintenance of data easier.2.4.2.2 Cloud Storage: Another convenient way to handle images, and other binary objects, is to shift responsibility to cloud-based storage centers.

Storage-as-a-service provides features to easily scale computational resources and provide access permissions. It also offers a user interface, command-line tools, and an API for several programming languages.3 https://en.wikipedia.org/wiki/Cron 4 https://nifi.apache.org/ 5 https://kafka.apache.org/ 6 https://www.ibm.com/products/mq 7 The most popular database for modern apps | MongoDB 8 https://hadoop.apache.org/docs/r1.2.1/hdfs_design.html14222.4.2.3 Database Storage: Database storage is another possibility. Here, the data, including images, with the associated metadata can be stored in one schema, which is the biggest advantage of this approach. Nevertheless, this approach is more appropriate for a small number of images and for applications that need to search data in real-time.The decision of what kind of storage to use depends on the target use case.

If there are no obstacles in uploading images to cloud-managed storage and it’s possible to pay for the external software the choice should be cloud-managed storage. It is easier to maintain and there is no need to worry about space limitations. On the other hand, if data is private and should be kept in local systems then paid cloud storage may not be a good option. A file system storage may be a better fit in this case. However, it is one that requires more resources and skills to set up and later administrate. On the other hands, a file system leaves more freedom to customize and adjust the solution to specific target use cases.2.4.3 Partitioning: Regardless of the choice of local or cloud-based file storage, special attention should be paid to partitioning. Partitioning data is crucial because it can drastically affect data processing. By partitioning data, the amount of data scanned by each query is smaller, thus improving performance and reducing cost.

By partitioning data, the amount of data scanned by each query is smaller, thus improving performance and reducing cost. A common practice is to partition the data based on time, often leading to a multi-level, partitioning scheme; but many other attributes may be used.2.5 Downgrading data amount To achieve sustainability, the system needs to be capable of managing different data sizes and save only the data necessary for future use cases. Several policies need to be defined: deleting the old data, aggregating the data, removing duplicates, and reducing data quality if needed.2.6. Building Decision Models To gain value from the data, many possibilities exist. One is to create an AI-based, decision- making model and another is to build a rule-based, expert system. AI-based models include predictive and clustering models. Predictive models need a solid history of data labeled with the information that needs to be predicted in the future.

AI-based models include predictive and clustering models. Predictive models need a solid history of data labeled with the information that needs to be predicted in the future. Clustering models create a group of similar data items that help experts use to support decision making. Rule-based expert systems are built using if-then rules, which are defined on the top of the fields of the data instances. Rules are defined by the experts in the specified area. Common practice is to combine both AI-based and rule-based approaches to build reliable, decision-making systems.2.7. Using Decision Models in Practice There are several options on how decision-making models can be used in practice. In Additive Manufacturing, for example, if a model can predict a critical event, an alert can be sent to staff in charge of handling that issue. If the model reports an in-process, monitoring anomaly, it can be used to change process parameters or stop the build.3.

If the model reports an in-process, monitoring anomaly, it can be used to change process parameters or stop the build.3. A variety of commercially available sensor technologies are now used in monitoring and controlling AM laser-based powder bed fusion (LBPF) processes. These technologies can be classified into two types: local sensors and global sensors. Local sensors have smaller fields of observation and acquire data with finer spatial resolutions and/or higher sampling rates. Global sensors often monitor the whole build surface and acquire data at lower resolutions. The two types complement each other and are used together for monitoring AM processes and predicting build quality [2].In-process Data Integration for Additive Manufacturing1423Camera-based, melt-pool monitoring (MPM) involves local sensing which generates high- resolution, melt-pool images at a high sampling rate.

MPM systems are critical for deep process understanding, in-process defect detection, and real-time control. Hence, they are one of the most attractive solutions for LPBF in-process monitoring [3]. However, MPM data streaming and integration for both real-time processing and offline analysis is very challenging for two reasons: the high velocity of the data generations and the big volume for archiving. In the following sections, the proposed data-integration framework is applied to MPM to illustrate and validate its benefits.3.1 Applying the Framework to AM 3.1.1 Requirements: To implement the framework for this use case, the requirements are first collected. The main requirements include knowledge of the input data, the capabilities of the system, the availability of storage, and standards for data flows. According to the framework, we obtained data source definition.

Based on that, an application of the framework will proceed differently. The last prerequisite is to provide storage where the data will be stored. Since the framework proposes separate storage for metadata and images, it is necessary to provide both.3.1.2 The Architecture: Figure 2 shows the architecture we propose for AM MPM data integration. The following sections describe how we use the framework with this architecture.Data Queue (Kafka)Figure 2. AM in-process Data Integration ArchitectureCollecting Data: Data is collected using the Camera link protocol. The Camera Link Driver is responsible for collecting data from the sensors and forwarding them to the queue.Queueing: The key to implementing this architecture is to have a buffer that will prevent the system overload, since the large amount of data must be processed in a short period. Apache Kafka10, an open-source message queue, was successfully used for this purpose in our set up.

Apache Kafka10, an open-source message queue, was successfully used for this purpose in our set up. Kafka can process many messages with negligible latency [8]. Also, Kafka is distributed, scalable system that offers high availability, so there is virtually no possibility of losing data.Raw data from the camera-link driver is stored in a queue. To ensure that the system can work with the amount of data expected, it must be ensured, among other things, that the Kafka cluster can withstand that amount of data. The buffer can be considered as a single point of failure, which means that in case it does not work, the system will not respond. Therefore, it is of great importance that there is a reliable and redundant component in place such that the buffer is highly unlikely to fail.10 http://kafka.apache.org/1425To ensure that Kafka can support the expected data flow, an overview of the capacity estimation is given below.

In this use case, we assume that the smallest interval of change in an image is 50 microseconds. Each frame of the camera is 120 by 120 in grayscale; hence, the image file size is approximately 15KB. Considering that, 20000 * 15KB = 300 MB of traffic is expected in one second. LinkedIn engineering has published a benchmark for a three-node Kafka cluster with a simple configuration on servers with 7200 RPM SATA drives and 32 GB RAM, which are connected via 1Gb Ethernet. The benchmark confirms that traffic of 800K records per second can be processed by a cluster configured in this way without any anomalies in performance. The process was monitored until the traffic grew to 1400GB [7]. In the use case presented in this paper, the same server configuration was used. The data would be stored in a buffer for the next 24 hours from the time it is received to ensure easy system recovery in case of overload of any of the following components.

The data would be stored in a buffer for the next 24 hours from the time it is received to ensure easy system recovery in case of overload of any of the following components. Considering one hour of high traffic, the calculation is 300MB * 60 seconds * 60 minutes ~ 1100 GB, which is within the range that a cluster configured in this way can process, based on the LinkedIn's benchmark (1400 GB). During longer use of this system, the cluster can be further configured to be even more suitable for this use case.From the queue, the images are sent for real-time analysis and for long-term archiving in parallel. We used Data Transfer, a custom application for preprocessing the data, for processing images and their metadata. In this way, we avoid storing irrelevant data in persistent storage.Archiving: Different types of images are involved in this use case. HDFS, the file system of the Apache Hadoop, was used for storing them.

HDFS, the file system of the Apache Hadoop, was used for storing them. Their metadata, however, were stored in Additive Manufacturing Materials Database (AMMD)11, built on top of MongoDB. HDFS is responsible for storing large, structured, and unstructured datasets across various nodes. Data arrives frequently so it is not advisable to archive it in real-time. That could overwhelm the HDFS, because it is not designed to work well with many small files. For that reason, Sequence files are used that merge many images in one file. Approximately, 8739 images should be merged so the size of one sequence file corresponds to the size of the minimal HDFS block, which is 128MB. In this way, the HDFS block will be fully used. Also, queuing from the previous step will create a time delay to transfer the data to permanent storage less frequently and avoid an overload.

In this way, the HDFS block will be fully used. Also, queuing from the previous step will create a time delay to transfer the data to permanent storage less frequently and avoid an overload. The authors in [6] suggested MapFiles as another solution that could be used for storing images, but further research is needed before using this in real use cases.AMMD is built using the NIST Material Data Curation System (MDCS) as a backend with structure provided by NIST’s AM schema. Providing a collaboration platform, AMMD is set to evolve through open data access and material data sharing among the AM community.Real-time Analysis: Next to being archived for future use, images and corresponding metadata are also used for real-time or near real-time analysis. Real-time functions include measurement data preprocessing - cleaning, melt pool feature extraction, anomaly detection, real-time feedback control generation, and emergency reaction, such as build stop.

Near real-time functions fuse the melt pool monitoring data obtained from previous layers, conduct layer-wise process and part state evaluation, and make decision if the build should be continued or stopped. The real-time and near real-time analysis generate events, alerts and alarms which are sent to the Alarm Management system for operator to perform reasoning and take actions. An automated alarm management system can be built to replace humans in conducting the cognitive task by querying additional data from other systems, for example, the metadata data store, MES or ERP system.11 https://ammd.nist.gov/14263.1.3 Data Flows: In this architecture, there are several data flows that need to be defined.

They are (1) between sensor and driver; (2) driver and Kafka; (3) Kafka and Data Transfer; (4) Data Transfer and Big Data Hadoop (5) Data Transfer and AMMD; (6) Kafka and feature extraction; (7) feature extraction and automated alarm management; (8) AMMD and automated alarm management; (9) AMMD and Big Data user.For 7-9, we propose OAGI Message models or OPC UA standards, while the question remains which standards should be used for the 1-6 data flows. This requires further research because it is necessary to consider the technologies that have been proposed and their possibilities. Figure 3 shows a proposed high-level OAGI message for Data Flow marked with 9, in a typical BOD structure. A new noun named “AdditiveManufacturingBigDataPackage” is created to capture both the metadata and the reference links to the big datasets [4].Figure 3.

A new noun named “AdditiveManufacturingBigDataPackage” is created to capture both the metadata and the reference links to the big datasets [4].Figure 3. Additive Manufacturing Big Data Package OAGi message4.DiscussionThe AM data integration presents a new application of the five-layer ISA 95 architecture12. Layer 0 involves the functions and standards associated with the physical production process. Layer 1 involves the functions and standards associated with sensing and manipulating that physical process. Layer 2 involves the functions and standards associated with automatically monitoring and controlling that process. Layer 3 has manufacturing operation functions and Layer 4 are enterprise functions.12 https://www.isa.org/standards-and-publications/isa-standards/isa-standards-committees/isa9514274.1 Architecture AlignmentAM data integration involves data flows for both monitoring and control and streaming for big data archiving.

Yet, we are mapping our architecture in Figure 2 to the lowest layers of the ISA architecture.Alarm management provides Layer 2 functions we use for “controlling the state of the AM process”. In addition, data integration for alarm management involves 1) the functions and standards associated with the physical production equipment (Layer 0) and 2) the functions and standards associated with sensing and manipulating that physical equipment (Layer 1). So, our AM data integration is functionally aligned with the ISA architecture for the branch of monitoring and control.At the same time, data flows are no longer restricted - in both time and space - to adjacent, functional layers defined in ISA 95. For example, in-process data can be directly streamed and integrated with Layer 3 and Layer 4 functions. This data then becomes available to be analyzed and used for real-time decision making as well.

This data then becomes available to be analyzed and used for real-time decision making as well. Together with other enterprise data, proactive actions can be taken to optimize enterprise performance and improve the manufacturing strategies. Even for the lower-level integration, the interfaces will be different. They will contain new types of information and they will require new types of data structures; or, develop new ways of using or extending existing structures. We chose the latter. On-going discussions with both OAGi and MTConnect are focused on extending their message structures, creating new ones, and using/sending AM data in that resulting structure. The discussions identified a list of required messages, their content, and their senders/receivers.

The discussions identified a list of required messages, their content, and their senders/receivers. Technical discussions were conducted with a focus on finding technical approaches and solutions to meet those requirements.4.2 Information Exchange StandardsRepresentative standards need to enable information integration and software interoperability both within and across each layer in Figure 2. Today, MTConnect and OPC UA13 are the two major standards adopted by AM machine vendors for machine data integration. The MTConnect standard offers a semantic vocabulary and data models to grab data from machine tools, production equipment and other factory hardware. Data from shop floor devices is presented in XML format, and is retrieved from information providers, called Agents, using HTTP as the underlying transport protocol.

Data from shop floor devices is presented in XML format, and is retrieved from information providers, called Agents, using HTTP as the underlying transport protocol. OPC UA is a two-way communication protocol enabling either sending request and response messages between clients and servers or a publish-subscribe model.In addition, a wide variety of Internet-of-Things connection options are emerging which might be applied to AM in-process data integration [5]. For instance, MQTT is a publish/subscribe protocol with minimal overhead and reliable communications, good for supervisory control and data acquisition (SCADA) and remote networks. Constrained Application Protocol (CoAP) provides the interoperability of HTTP but with minimal overhead, an appropriate substitution for edge- based devices where HTTP would be too resource intensive.

Digital Data Service (DDS) is an open publish/subscribe protocol for fast and decentralized communication, best for machine-to- machine (M2M) communications.13 https://opcfoundation.org/about/opc-technologies/opc-ua/1428All the communication protocols mentioned above provide an effective mechanism to integrate small volume data sampled at rates needed in traditional manufacturing processes. For the high velocity, high volume data generated from field devices, new standards are needed to integrate new types of data such as images. Given the complexity of the architecture and its new technology components required for AM data integration shown in Figure 2, existing standards must be re- evaluated and improved for smart AM.5.SummaryThis paper offers a comprehensive framework for data integration in additive manufacturing and, moreover, shows the practical implementation of the framework on the example of a melt pool monitoring system.

In the paper, the key components of the framework are highlighted and an overview of the main decisions that must be made for the proper use of each step of the framework is given. Moreover, the proposed framework can be applied to other additive manufacturing systems and quickly achieve the desired results whether the goal is to archive data for future use or real-time analytics. The paper points that, while standards for the higher-level processing data flows are being developed already, there is a current lack of standards for messages exchanged in preprocessing data flows. Further research should include defining messaging standards for each of the observed flows.DisclaimerCertain commercial systems are identified in this paper. Such identification does not imply recommendation or endorsement by NIST; nor does it imply that the products identified are necessarily the best available for the purpose.

Such identification does not imply recommendation or endorsement by NIST; nor does it imply that the products identified are necessarily the best available for the purpose. Further, any opinions, findings, conclusions, or recommendations expressed in this material are those of the authors and do not necessarily reflect the views of NIST or any other supporting U.S. government or corporate organizations.References1. K. Morris, Y. Lu, and S. Frechette, "Foundations of Information Governance for Smart Manufacturing," Smart and Sustainable Manufacturing Systems 4, no. 2 (2020): 43-61.2. P. Boulware, “In-Process Monitoring Techniques for Laser Powder Bed Fusion”, https://ewi.org/wp-content/uploads/2018/12/LP4-Boulware_In-Process-Monitoring-Techniques- for-Laser-Powder-Bed-Fusion.pdf, Accessed on June 28, 20213. Y. Lu, Z. Yang and J.

viewed PaperA Data-Driven Approach for Multi-Lattice Transitions Martha Baldwin,1 Nicholas A. Meisel,2 Christopher McComb11Department of Mechanical Engineering, Carnegie Mellon University, Pittsburgh, PA 15213 2School of Engineering Design, Technology, and Professional Programs, The Pennsylvania State University, University Park, PA 16802AbstractAdditive manufacturing is advantageous for producing lightweight components while maintaining function and form. This ability has been bolstered by the introduction of unit lattice cells and the gradation of those cells. In cases where loading varies throughout a part, it may be necessary to use multiple lattice cell types, also known as multi-lattice structures. In such structures, abrupt transitions between geometries may cause stress concentrations, making the boundary a primary failure point; thus, transition regions should be created between each lattice cell type.

Although computational approaches have been proposed, smooth transition regions are still difficult to intuit and design, especially between lattices of drastically different geometries. This work demonstrates and assesses a method for using variational autoencoders to automate the creation of transitional lattice cells. In particular, the work focuses on identifying the relationships that exist within the latent space produced by the variational autoencoder. Through computational experimentation, it was found that the smoothness of transition regions was higher when the endpoints were located closer together in the latent space.1. IntroductionAdditive manufacturing (AM) is renowned for producing lightweight components while addressing complex geometry and strength requirements. This ability has been bolstered by the introduction of unit lattice cells, which have enabled designers to significantly reduce the weight of a part, while maintaining needed stiffness and geometry[1,2].

Proposed applications of lattices are often centered around light weighting or improving the stiffness to weight ratios of components manufactured [1]. Lattices are also discussed as impact reduction applications, such as the crumple zone in a car [1] or surgical implants that experience variable loading [3]. Recent work has indicated that graded lattice structures are the next step for improving upon uniform lattice structures [4–8]. These structures vary the thickness of individual strut-based unit cells to accommodate areas of high or low stress. Graded lattice structures have been shown to significantly outperform the stiffness of uniform lattice structures of comparable weight [5,6,8]. Given the unique deformation properties of graded lattices, it is believed that they would be more beneficial than uniform lattices in impact reduction applications [9].

Given the unique deformation properties of graded lattices, it is believed that they would be more beneficial than uniform lattices in impact reduction applications [9]. Further research has explored the possible improvements afforded by multi-lattice structures, which are structures comprised of multiple types of unit cell topologies [3,10–13]. Similar to graded lattice structures, multi-lattice structures have higher strength and stiffness than uniform lattice structures of comparable density[9,12]. However, in multi-lattice structures, abrupt transitions between lattice types may cause stress concentrations, which are detrimental to overall part strength [12]. One method to address abrupt transition regions is to interpolate between the faces of two lattice cell types [10]. This creates a smooth transition which evenly distributes the stresses throughout the structure.

This creates a smooth transition which evenly distributes the stresses throughout the structure. However, smooth transition regions are currently difficult to intuit and design, especially between lattices of drastically different geometries.2197Recent work has utilized machine learning to create interpolations between unit cells by optimizing the properties of the cells [14,15]. While optimization-based methods are powerful, they leave little space for input from the user. This work explores the effectiveness of convolutional Variational Autoencoders (VAEs) for automating the creation of transition regions in multi-lattice structures. Specifically, we design a VAE system that can predict transitions between different unit lattice cells and produce transition regions with any length. Our primary research questions address the unknowns about the effects of dimensionality reduction with respect to lattice topologies: 1.

Our primary research questions address the unknowns about the effects of dimensionality reduction with respect to lattice topologies: 1. How does distance in the latent space and transition length affect the smoothness ofinterpolations?2. How do differences in unit cell topology affect smoothness?2. Background 2.1 Unit Cells for DesignThis section defines lattice unit cells and their usefulness to the research and design community, as well as establishing the motivation for using multi-lattice structures. In this paper, the focus will be on periodic lattice structures, which are created by repeating a unit cell. A uniform lattice structure consists of one type of unit cell, where all the cells in the structure have the same density and size. The stress strain curve of uniform lattices matches standard plastic deformation, where there is an elastic region, followed by a plastic plateau, until the object reaches full densification [7].

The stress strain curve of uniform lattices matches standard plastic deformation, where there is an elastic region, followed by a plastic plateau, until the object reaches full densification [7]. The properties of each type of unit lattice are unique [1], making it important to select the correct type of lattice for a design. A graded lattice structure consists of one type of unit cell, but the cells in the structure can vary in volume fraction [6]. It has been proven that functionally graded lattice structures perform better with respect to stiffness [5,6,8]. The benefits of graded structures are the additional design freedom that can be used to customize failure modes and stress-strain responses [7,8]. It was suggested that such properties would be beneficial in applications where there are dynamic loads, such as surgical implants [3] or impact protection equipment [9]. These structures are also relatively simple to create. Li et al.

These structures are also relatively simple to create. Li et al. created graded structures by using relative density mapping from a topology optimization and assigning densities based on stress requirements [4]. However, the key restriction in graded lattice structures is their property dependence on the type of unit cell. It was found that surface-based unit cells outperformed strut-based cells due to their connectivity[8], as well as the degree of gradation greatly affecting the cumulative energy absorption for body-centered cubic lattices, but not Schwarz-P lattices [2].The term “multi-lattice” structure is being defined as a structure that contains multiple types of unit cells. This term is being proposed in this work to differentiate from the other types of mesostructures that are possible. The motivation for this work revolves around the implementation of multi-lattice structures where the transitions between unit cells are continuous.

The motivation for this work revolves around the implementation of multi-lattice structures where the transitions between unit cells are continuous. Many studies that explore multi-lattice structures have utilized unit cells with matching boundaries to avoid developing a method to connect nodes [3,9,12,13]. Although this is not an area of interest, it is important for understanding the behavior of multi-lattice structures on a rudimentary level. To create the structures, a relative density mapping from a topology optimization was executed to assign 2.5-dimensional unit cells based on the stresses in the system [9,12,13]. As a result, they were able to prove that trusses utilizing 2.5-dimensional multi-lattice structures outperform uniform lattice structures regarding stiffness and strength[9,12,13]. Another study was conducted by Gok that proved that using a multi-lattice hip implant reduced the maximum stress and weight2198compared to conventional hip implants [3].

Another study was conducted by Gok that proved that using a multi-lattice hip implant reduced the maximum stress and weight2198compared to conventional hip implants [3]. Although the work does not provide comparison between other types of mesostructured patterning, it does demonstrate a strong use case for multi- lattice structures. It should be noted that the study was also restricted to unit cells that had overlapping boundaries, so the transition regions were inherently smooth.Although multi-lattice structures are more beneficial, there are some potential drawbacks. Kang et al. concluded that the boundaries between unit cell types caused stress concentrations [12]. Which indicates that there is a need for smooth transitions between unit cells to appropriately distribute stress through structures. Sanders et al. created transition regions by using signed distance functions with respect to the boundaries of each strut [10].

Sanders et al. created transition regions by using signed distance functions with respect to the boundaries of each strut [10]. By combining transition regions with functional grading, they are able to prove the effectiveness of multi-lattice structures. Their work studies topology optimization of multi-lattice structures that use functionally graded interpolated transition regions. Although this technique can be applied to “unit cells composed of noncylindrical bars or plates”, the interpolations between each pair of unit cells must be computed individually. For those looking to optimize both the macrostructure and the mesostructure, this method is extremely repetitive and tedious. Wang et al.

For those looking to optimize both the macrostructure and the mesostructure, this method is extremely repetitive and tedious. Wang et al. have also regarded the creation of transition regions as “a challenging problem involving complex inverse design at the microscale, costly nested optimization at the macroscale, and boundary matching between neighboring microstructures [14].” Which is why more recent research has explored how to create smooth transitions between unit cells using machine learning methods [10].2.2 Machine Learning MethodsRecent approaches to creating continuous transitions in multi-lattice structures have utilized machine learning.

This section establishes the two notable machine learning methods which have been used for creating multi-lattice structures: Variational Autoencoders (VAEs) and Generative Adversarial Networks (GANS).Generative Adversarial NetworksGenerative Adversarial Networks (GANs) utilize two models that work in parallel to ultimately create a single model for generating data [16]. These two models are known as the generator and discriminator, which work against one another in effort to beat the other. The generator produces fake data that is learned from real data, while the discriminator learns to sort the fake and real data. As training progresses, each model progressively becomes better at its purpose until the discriminator is unable to distinguish between real and fake data produced by the generator.Wang et al. used an Inverse Homogenization GAN to generate a multi-lattice structure that reduced stress concentrations by nearly 80% [15].

used an Inverse Homogenization GAN to generate a multi-lattice structure that reduced stress concentrations by nearly 80% [15]. This method used a topology optimization of the design space to create a relative density mapping, while considering the type of unit cells being used. This serves as the main motivation for this work, as it proves there is a benefit to creating multi-lattice structures. The advantage to using a GAN, was that it could restrict designs based on manufacturability [15].However, this work was limited by the type of unit cells that could be interpolated. The focus was on triply periodic minimal surfaces, as their physical characteristics could be described by 6 principal components. Indicating that only structures generated using functions would be suitable for this method, which limits the freedom for using personalized unit cells.

Indicating that only structures generated using functions would be suitable for this method, which limits the freedom for using personalized unit cells. The primary2199drawback to this method was that it was only intended to organize the types of unit cells, but only based on performance rather than shape. Therefore, interfaces between the cells were not perfect, and a second interpolation had to be performed to create smooth transition regions. Indicating that the IH-GAN was not ideal for creating interpolations, but better suited for optimization of unit cell types.Variational AutoencodersA variational autoencoder (VAE) is a machine learning model that learns how to perform data-driven dimensionality reduction [17]. Reducing the dimensionality of data can make it more computationally efficient to perform inferences, such as interpolations.

Reducing the dimensionality of data can make it more computationally efficient to perform inferences, such as interpolations. What makes VAEs unique from other techniques is that they perform a non-linear dimensionality reduction, which is advantageous for reducing complex data types [18]. The reduced data is combined to establish the reduced dimensionality latent space, which can be used to represent complex mechanical parts [19] and performance characteristics of engineered systems [20].A VAE consists of two models, an encoder, and a decoder, which work in series. Wherethe encoder acts as a recognition model and the decoder as the generative model [16,18]. The encoder performs a dimensionality reduction which creates a latent representation of the data. The decoder uses the information from the latent space to reconstruct the data. Finally, the VAE is trained by minimizing the error between the decoder model and the original data provided.Wang et al.

Finally, the VAE is trained by minimizing the error between the decoder model and the original data provided.Wang et al. constructed a VAE to perform multi-lattice interpolations between 2D lattices [14]. This work created interpolations by traveling along the shortest paths of trained points in the latent space. Another application of VAEs in AM has been describing the properties of randomly generated unit cells [18]. Xue et al. used a principal component analysis (PCA) reduction of their generated 2D unit cells to classify them based on physical performance [18]. Their work proved that location in the latent space correlated with young’s modulus, and shear modulus. Given the nature of lattice topologies, VAEs are a good candidate for dimensionality reduction of these structures. Wang et al.

Given the nature of lattice topologies, VAEs are a good candidate for dimensionality reduction of these structures. Wang et al. identify 3 benefits to the latent space created by VAEs [14]: interpolations can be created by traveling in the latent space, shape similarity can be evaluated based on distance in the latent space, and the latent space can be used to map the physical properties of the unit cells.In contrast to existing work, our focus is to establish a methodology that permits a designer- centered approach rather than a fully automated VAE optimization. Embracing a human-in-the- loop paradigm permits the designer to inject stylistic elements as well as easily account for aspects that would be difficult to define numerically. This interaction will currently consist of selecting the endpoint unit cells and the length of the transition region.3.

This interaction will currently consist of selecting the endpoint unit cells and the length of the transition region.3. MethodologyIn this section, the process of data generation, development of the VAE model, and execution and evaluation of interpolations are detailed. The current work specifically deals with interpolation of 2-dimensional lattice unit cells in order to simplify the data needs and training time of the VAE model. The methodological approach to testing the hypotheses of our research questions is visualized in Figure 1. Section 1 will describe how artificial data was generated to use for testing. Section 2 outlines the hyperparameters and key characteristics of the VAE used for2200training. Section 3 shows the process for creating interpolations in the latent space.

Section 2 outlines the hyperparameters and key characteristics of the VAE used for2200training. Section 3 shows the process for creating interpolations in the latent space. Section 4 will develop the smoothness metric to measure the performance of the interpolations.Figure 1: Illustration of overall methodologies3.1 Synthetic Data Generation To begin training the VAE, a significant amount of data was necessary to produce a sufficiently dense latent space for interpolation. Synthetic data was generated based on a series of 12 shapes (Figure 2) that mimicked other additive manufacturing literature[9,13]. Although the dataset does not contain a wide variety of shapes, it will allow for thorough testing of the latent space itself, as our focus is on how the quality of interpolations change within the latent space.Figure 2: Synthetic Shape TypesTo add variation to the data, the thickness of the shapes and densities were varied.

Where the thickness of the shapes represents the increasing size of the struts. This was done stepwise by expanding upon the existing shapes (See top of Figure 3). The density represents the value of each pixel between 0 and 1, which are black and white respectively. This representation of density is indicative of the likelihood of a pixel being present (See bottom of Figure 3). Once all the shapes were generated, there were a total of 415 data points. Although limited, this dataset proved sufficient for the purposes of this paper.2201Figure 3: Original shape with possible density and pixel intensity variations3.2 Training the Variational Autoencoder This section outlines how the VAE was designed for this research. When the VAE was created, there were many factors that had to be considered, as there are many hyperparameters of VAEs that affect their performance. First, the overall framework of the VAE affects the performance and outputs from the system (See Figure 4).

First, the overall framework of the VAE affects the performance and outputs from the system (See Figure 4). The framework shown was designed to mimic the VAEs used to organize images. Another important hyperparameter is the latent dimensionality. In this case, the latent dimensionality was chosen based on the qualitative difference between the original and decoded unit cells. A qualitative study of latent space dimensionality revealed that a dimensionality of 4 provided a desirable balance between training time and reconstruction accuracy.Figure 4: Framework of VAE1There are many additional features that can be incorporated into a VAE to improve performance. In this case, training was terminated early if the loss failed to improve after several epochs and the autoencoder was also designed to save the weights from the best iteration. The system used a batch size of 32 to and the adam optimizer.

The system used a batch size of 32 to and the adam optimizer. Given this framework, the VAE used 85% of the dataset to train and 15% to validate and test the performance of the VAE.3.3 Creating InterpolationsOnce trained, the autoencoder has established a latent space in which interpolations can be performed. The latent space represents a combination of the dimensionally reduced training data. The encoder serves to recognize the unit cell and assign it a predicted latent point, and the decoder then recreates the unit cell using the predicted latent point. These functions would be used for creating interpolations between the different unit cells. As mentioned previously, interpolations require an encoding and decoding to produce desired results.

These functions would be used for creating interpolations between the different unit cells. As mentioned previously, interpolations require an encoding and decoding to produce desired results. The unit cells at the end of the desired1 Generated with this site: https://dropsofai.com/variational-autoencoders-and-image-generation-with-keras/2202transition region are first provided to the encoder, which calculates the latent points of the two cells. The transitions points are then interpolated between the encoded endpoints in the latent space.

The transitions points are then interpolated between the encoded endpoints in the latent space. Finally, the decoder generates the cells that correspond to those latent points (See Figure 5).Figure 5: Diagram demonstrating interpolation process3.4 Evaluating Smoothness To accurately gauge the quality of an interpolation, a metric was designed to measure the smoothness of the transition region, where smoothness needed to (1) measure how the interpolation changed both within images and across images and (2) penalize pixels disconnected from the main structure, as we did not want to encourage the generation of floating pixels. Finally, (3) the metric should account for pixel intensity, allowing it to account for the gradation of pixels, as interpolations will often produce non-binary images. Related metrics have been published in the literature [15,21–24], but each of these metrics fails to satisfy at least one of the aforementioned criteria.

Related metrics have been published in the literature [15,21–24], but each of these metrics fails to satisfy at least one of the aforementioned criteria. A custom measurement metric was designed to emphasize change along the edges of the unit cells using a 3-dimensional Sobel filter [25,26]. This makes it possible to resolve both within image changes (the x and y directions shown in Figure 6) and between-image changes (the z direction shown in Figure 6) using a single operation.Figure 6: Visualization of the model convolved with Sobel filterSpecifically, we use the formulation of 3D Sobel operators defined by Amin et al.

Then the root mean squared error (RMSE) between the consecutive gradients was measured, as that would be the indicator of “smoothness” between each dimension (Eq 3).𝑅𝑀𝑆𝐸!,% = 	 7)+’∑ 9𝐺!,%&’,) − 𝐺!,%,): 𝑁(𝑅𝑀𝑆𝐸",% = 	 7( ∑ 9𝐺",%&’,) − 𝐺",%,): 𝑁)+’𝑅𝑀𝑆𝐸#,% = 	 7( ∑ 9𝐺#,%&’,) − 𝐺#,%,): 𝑁)+’where 𝑅𝑀𝑆𝐸!, 𝑅𝑀𝑆𝐸", 𝑎𝑛𝑑	𝑅𝑀𝑆𝐸# are the root mean squared errors of a pair of gradients in the x, y, and z directions respectively, j is the index that identifies the specific term in the gradient array, and N is the number of terms in a single gradient array.Once the RMSE was calculated for the x, y, and z components, the average RMSE wascalculated and normalized to create a final “smoothness value”.𝑅𝑀𝑆𝐸% =𝑅𝑀𝑆𝐸!,% + 	𝑅𝑀𝑆𝐸",% + 𝑅𝑀𝑆𝐸#,% 3 ⋅ 𝑅𝑀𝑆𝐸,-!where 𝑅𝑀𝑆𝐸,-! is the maximum possible 𝑅𝑀𝑆𝐸% which is calculated based on the filter used.

The normalization of the RMSE allowed for the smoothness value to be represented as a percentage.2204( 1 )( 2 )( 3 )( 4 )𝑠𝑚𝑜𝑜𝑡ℎ𝑛𝑒𝑠𝑠 = (1 − 𝑎𝑣𝑔(𝑅𝑀𝑆𝐸%)) ⋅ 100( 5 )To validate and determine the limits of this measurement, multiple baseline evaluationswere conducted. These evaluations consisted of the following (Table 2):1. 1-pixel step-wise interpolation with decreasing pixel intensity values 2. Row-step-wise interpolation with transition occurring at fully solid shape 3. A “smooth” interpolation with abrupt changes 4. Random pixel arraysEvaluation 1 was to determine the best possible interpolation value, as it is impossible to achieve 100% smoothness unless the unit cells are identical. Evaluation 2 was simply a test interpolation to determine the performance of the metric if the pixels were purely black and white. Evaluation 3 was to determine how the interpolations would be affected if a “bad” transition step was introduced in random locations.

Evaluation 3 was to determine how the interpolations would be affected if a “bad” transition step was introduced in random locations. Evaluation 4 was chosen to evaluate the metric when the transitions lack a clearly defined edge.Table 1: Smoothness baseline metric evaluationsInterpolationSmoothness98.797%80.5%66.775%79.879%4. ResultsThis section reports the performance of the trained VAE, assesses the qualitative and quantitative results from interpolations, and addresses the research questions outlined in the introduction. Overall, the purpose of this experiment was to demonstrate the feasibility of using VAE's to interpolate 2D unit cells and determine a metric for evaluating such interpolations.4.1 Performance of Convolutional Variational Autoencoder (VAE)Before using the results from a VAE, it should be determined whether the VAE is overfitting or underfitting the data.

The performance of the VAE was measured using loss and coefficient of determination (R2 value), which can be seen in Figure 7. The R2 value was in excess of 95%, which indicates that the model is a good fit of the training data. The significant decrease in loss indicates that the generated unit cells began to match the original unit cells more closely. Given these results, the trained VAE was considered satisfactory for testing.2205Figure 7: Performance of VAE with a latent dimensionality of 44.2 Interpolation ResultsThis section demonstrates how the smoothness evaluation is executed, and what it outputs. To address the research questions in the introduction, this section also shows the testing of different types of interpolations. Note that the decoded endpoint unit cells do not perfectly match the user provided unit cells (See Figure 8(a)), this was due to use of the test unit cells for interpolation.

Note that the decoded endpoint unit cells do not perfectly match the user provided unit cells (See Figure 8(a)), this was due to use of the test unit cells for interpolation. Since the VAE is trained on 85% of the data, it created the best reconstructions possible to match the provided test points. It was expected that the encoder would not be capable of perfect recreations of test points since the VAE contains loss. Whereas Figure 8(b) shows a method for visualizing the latent space, while helping to show where the interpolation(a) is oriented with respect to the rest of the latent space.2206Figure 8: Demonstration of Interpolation: (a) Interpolation (b) PCA reduced latent space plot with the interpolation points super imposed in the spaceHow does distance in the latent space and transition length affect smoothness? It was hypothesized that the distance in the latent space would result in a decrease in smoothness as points would become spread further apart.

Where the length of the transition region would have a proportional relationship with the smoothness of the interpolations. To test this hypothesis, a variety of different interpolations were defined and assessed. First, the mean and standard deviation of the latent space were calculated, since the latent space is unitless, the standard deviations within the space serve as the best measurement for comparison between trained VAEs. Interpolations were performed from -3 standard deviations to 3 standard deviations from the mean. We tested sets of interpolations consisting of 5, 10 and 15 transition points.

Interpolations were performed from -3 standard deviations to 3 standard deviations from the mean. We tested sets of interpolations consisting of 5, 10 and 15 transition points. A sample of the interpolations can be seen in Table 2 for transition length 10.2207Table 2: Visualization of transitions based on number of standard deviations with 10 interpolation pointsNumber of Standard Deviations Between EndpointsResulting Interpolations123456Visually, the unit cells are most similar when there are more interpolation points and 1 standard deviation between the endpoints. To confirm this, the smoothness evaluation was conducted on all interpolations (See Figure 9).2208Figure 9: Smoothness vs Number of Standard Deviations in the Latent Space. (a) Analysis based on 15 interpolation points - More interpolation points in the latent space result in slightly poorer transitions as distance increases.

(a) Analysis based on 15 interpolation points - More interpolation points in the latent space result in slightly poorer transitions as distance increases. (b) Analysis based on 10 interpolation points (c) Analysis based on 5 interpolation points - Fewer interpolation points result in significantly rougher transitions over a longer distance in the latent space.To fully confirm the results, an ordinary least squares (OLS) regression was executed on the data gathered in Figure 9, where the dependent variable was percent smoothness, and the independent variables were transition length and distance in the latent space. This regression was found to be highly statistically significant (p < 0.001) as well as practically significant (R2 = 0.924). The p-values in Table 4 confirmed that as the distance between endpoints increases, the smoothness decreases. The distance in the latent space has the most significant effect on smoothness.

The p-values in Table 4 confirmed that as the distance between endpoints increases, the smoothness decreases. The distance in the latent space has the most significant effect on smoothness. However, the results also indicate that transition length alone is unlikely to influence the smoothness value. Consequently, the combination of the two variables, noted by the cross-term in Table 4, does directly affect the smoothness. This indicates that the transition length only affects the smoothness when distance is also accounted for.Table 3: Ordinary Least Square RegressionConstant: Number of Standard Deviations (Distance): Transition Length: Cross-Term:Coefficient108.6807 -8.8616 -0.4182 0.4601Standard Error 3.882 0.997 0.359 0.092pp < 0.0001 p < 0.0001 0.264 p < 0.0001Given these results, the best interpolations can be produced by reducing the distance traveled in the latent space as much as possible.

In cases where the distance is vast between endpoints, increasing the number of transitions will only improve the smoothness by a small amount. The smoothness is heavily reliant on distance within the latent space.How do differences in unit cell topologies affect smoothness?This section will explore the effects of the qualitative differences in unit cell shape with respect to smoothness. It is hypothesized that as the endpoints of an interpolation become more “different”, their smoothness should decrease. Given the complexity of the latent space, visual2209inspection may be deceptive, which is why it is important to address the unexpected. Table 5 consists of an array of various interpolations, all of which have the same initial interpolation point. This was to ensure that the interpolations were directly comparable to one another.

Table 5 consists of an array of various interpolations, all of which have the same initial interpolation point. This was to ensure that the interpolations were directly comparable to one another. Although the distances between the endpoints are not the same, this series of interpolations are ordered based on smoothness performance, where the best interpolation is Example 1.Table 4: Examples of different topologies and their effect on smoothness of the interpolationExampleExample InterpolationsSmoothness (%)195.946295.398391.576490.928588.141The results from Table 5 were somewhat as expected, but this section will outline why the results here were reasonable. From a qualitative standpoint, the rankings were expected to order Example 1 > 2 > 4 > 5 > 3, however, the smoothness metric is based on the variation along all the unit cells, not solely the endpoints.

Example interpolations 1 and 2 performed similarly well, as both were a rotation of the initial endpoint. Example 3 performed slightly worse than 1 and 2, but there were a lot of pixels that needed to be activated. However, based on the qualitative evaluation it would appear that 3 should have performed the worst. Since the smoothness evaluation is highly dependent on the amount of similarity between consecutive unit cells, the endpoint unit cells in Example 3 would have increased the smoothness percentage significantly. Finally, example 4 performed slightly better than example 5, as the diagonal was not removed in example 4, meaning that fewer changes occurred.22105. DiscussionThis section will discuss the results obtained from a transition region developed between four different types of unit cells, while highlighting the many possibilities that are afforded by VAEs and their future work. This work underscores the potential of VAEs for supporting transition region design.

This work underscores the potential of VAEs for supporting transition region design. However, the smoothness of the transition region is primarily predicted by the distance in the latent space. Therefore, our hypothesis is partially refuted, as the transition length did not have the expected impact on smoothness.Figure 10(a) shows how powerful VAEs are by demonstrating a smooth transition region between four distinct unit cells. Such a structure represents how a transition boundary would be developed between four different lattice topologies. Multi-lattice designs incorporating more than two types of lattices could expand design possibilities but creating transition regions has been a main restriction. As far as we know, this type of transition has only been reproduced using a VAE, as it is extremely computationally intensive using other methods. Figure 10(b) represents the locations of each of the predicted points in the PCA reduced latent space.

Figure 10(b) represents the locations of each of the predicted points in the PCA reduced latent space. As established previously, the closer points will have smoother transitions based on the smoothness metric. This type of visual aid could be utilized to quickly determine which transition regions perform poorly.2211Figure 10: (a) Example of mesh grid interpolation (b) PCA reduced representation of the mesh grid interpolation in the latent space22126. ConclusionIn cases where the loading varies throughout a part, it may be necessary to use multiple lattice cell types, also known as multi-lattice structures. In such structures, abrupt transitions between lattice types may cause stress concentrations, making the boundary a primary failure point; thus, transition regions should be created between each lattice cell type. This work demonstrates and assesses a method for using variational autoencoders to automate the creation of transitions amongst multi-lattice structures.

This work demonstrates and assesses a method for using variational autoencoders to automate the creation of transitions amongst multi-lattice structures. In comparison to other computational approaches, this work focused on analyzing how a VAE latent space affected unit cell interpolations by answering two research questions.First, we answered how distance in the latent space and the number of transitions affects the smoothness of an interpolation. Specifically, we assessed the smoothness of lattice interpolations produced using a VAE, evaluated using 3D Sobel filters. Using an OLS regression, it was found that distance between endpoints in the latent space had the most significant impact on smoothness, whereas the number of steps in the transition were unlikely to affect the smoothness alone. Working in unison, the two variables only had some impact on the smoothness value.

Working in unison, the two variables only had some impact on the smoothness value. Given these results, the best interpolations can be produced by reducing the distance traveled in the latent space as much as possible. This indicates the need for using latent space optimization methods to minimize the distance between desired end points during training.Secondly, we qualitatively evaluated comparable interpolations to determine whether their performance met our expectations. The results from these tests indicate that human qualitative similarities may not perform as well as similarities made by the model. Based on the smoothness evaluations of our interpolations, humans cannot accurately and precisely determine which set of interpolations is most smooth.Future work will also focus on modifications to be made that encourage manufacturability of transition regions.

This will require that the connectivity between cells can be ensured, as the physical properties of the unit cells are dependent on their ability to be created. This will be followed by determining the best method for analyzing the functionality of these transition regions, as they are extremely complex structures which require physical analyses to be useful to designers. It is important to determine if the smoothness metric has any correlation to physical properties within a transition region. As discussed previously, other works have incorporated physical properties into their latent space optimization [14], which may prove to be the best option for organizing the latent space.In order to take advantage of the latent space created by VAEs, we must determine how to address the effects on distance within that space.

Given that the distance has such a huge impact on smoothness, the unit cells that perform best will exist in the center of the space where they are close to all the other unit cell. Unit cells on the boundaries of the latent space have less information, so they are difficult to decode without significant loss of the original unit cell. Furthermore, the resolution of the outputs is highly dependent on the resolution of the training data. The data size increases by an exponent of the data dimensionality. Consequentially, as this work expands into the 3-dimensional space, we will have to address extreme increases in data size and therefore training times of the VAE.2213References:[1]Hanks, B., Berthel, J., Frecker, M., and Simpson, T.

-GAIN LAYER-TO-LAYER STABILITY CRITERION IN LASER METAL DEPOSITION PROCESSESP.M. Sammons, D.A. Bristow, and R.G. LandersMechanical and Aerospace Engineering Department Missouri University of Science and Technology Rolla, Missouri 65409-0050AbstractIn Laser Metal Deposition (LMD), a blown powder metal additive manufacturing process, functional metal parts are fabricated in a layer-by-layer fashion. In addition to the in- layer dynamics, which describe how the process evolves within a given layer, the additive- fabrication property of LMD creates a second set of dynamics which describe how the process evolves from layer-to-layer. While these dynamics, termed layer-to-layer dynamics, are coupled with both the in-layer dynamics and the process operating conditions, they are not widely considered in the modeling, process planning, or process control of LMD operations.

Because of this, seemingly valid choices for process parameters can lead to part failure – a phenomenon commonly encountered when undergoing the laborious procedure of tuning a new LMD process. Here, a layer-to-layer stability condition for LMD fabrications is given, based on the shape of the powder catchment efficiency function, which provides insight into the layer-to-layer evolution of LMD processes and can be useful in process planning and control. The stability criterion is evaluated for various operating points, allowing stable and unstable operating regions to be identified. Simulation results are presented showing both stable and unstable layer-to-layer LMD fabrications.

Simulation results are presented showing both stable and unstable layer-to-layer LMD fabrications. The simulated behavior successfully predicts the results seen in both stable and unstable experimental depositions.IntroductionLaser Metal Deposition (LMD) is an additive manufacturing process in which a laser and a blown metal powder source are used to create functional metal parts in a layer by layer fashion. The LMD process, and additive manufacturing processes in general, are governed by coupled two-dimensional dynamics – the in-layer dynamics, which describe how the process evolves withing a given layer, and the layer-to-layer dynamics, which describe how the process evolves from a given layer to subsequent layers. However, LMD processes are typically modeled as purely in-layer processes and the layer-to-layer dynamics are neglected [1,2]. Neglecting the layer-to-layer dynamics in process planning and attempting to control the process can result in failed parts [3].

Neglecting the layer-to-layer dynamics in process planning and attempting to control the process can result in failed parts [3]. An example of a failed part is shown in Figure 1.Figure 1. Photograph of a failed constant process parameter deposition.Further, without a model of the layer-to-layer dynamics, insight into how the layer-to- layer dynamics cause part failure is impossible rendering process tuning and planning is extremely difficult.In order to gain insight into how the process evolves in both dimensions, a repetitive process model was developed in [4]. This model captures the dominant in-layer and layer-to- layer dynamics of the process while still retaining a simple structure allowing it to be used for process control design and planning.

This model captures the dominant in-layer and layer-to- layer dynamics of the process while still retaining a simple structure allowing it to be used for process control design and planning. While the developed model is dynamic in both the in-layer and layer-to-layer domains, knowledge of the process dynamics for particular systems [5] allows for the application of repetitive process control theory to the model to result in a simple expression to determine the layer-to-layer stability, based on the powder catchment efficiency function of the particular LMD process, for a given set of process parameters. The criterion enables the creation of simple, yet informative process maps for determining whether or not a particular set of LMD process parameters will yield a successful build.Model Description and PreliminariesLet hj denote the height profile of a deposition on a given layer j. On each layer, a new molten bead δhj is deposited on top of the existing height profile hj–1.

On each layer, a new molten bead δhj is deposited on top of the existing height profile hj–1. Mathematically, this height accumulation is described as an integration process in the layer domain,1jjjhhhwhere j is the layer number, h is the height of the part above the substrate (mm), and δh is the bead height, i.e., the incremental amount of material added to the structure (mm).

The bead height is given as10,1jSjjjhbfdhwhere b0 is the characteristic bead width (mm) and converts the area of the bead to the height of the bead, ς is the material specific volume (mm3/kg), fμ is the catchment efficiency function, dS is the distance between the substrate and the cladding head (mm) and a process input, λ = ṁ/v is the spatial powder flow rate (kg/mm) and the second process input, ṁ is the powder flow rate (kg/s), and v is the substrate scan speed (mm/s).During a typical operation, the two process inputs, spatial flow rate λ and substrate standoff distance dS, are commanded to be constant for a given layer. These two input signals, in addition to a set of initial conditions, result in a constant addition of material in the height(1)(2)direction during each layer.

These two input signals, in addition to a set of initial conditions, result in a constant addition of material in the height(1)(2)direction during each layer. The collection of inputs λr and dS,r, initial condition on the distance between the part and the cladding head dP,0, and outputs hr that satisfy (1) for all j is termed a solution trajectory. Because the solution trajectory describes a nominal operating process and (1) is a nonlinear discrete dynamic equation, the effects of perturbations to the process on the output are analyzed about this trajectory. Along the solution trajectory with , the first order linearization of (1) is,,SjSrdd1211jjjhhwhere κ1 is the derivative of f with respect to λ and κ2 is the derivative of f with respect to dS (or h),110,0Pbfd,02,002PcenterrscalePwidthdfdbjjrhhhjjr.

Equation (3) is the and the incremental variables are defined as linear DC-gain model of the LMD process presented in [4] and more details regarding the full derivation of the model, including the dynamics associated with the in-layer domain, can be found there. Of note is that (3) is a dynamic equation purely in the discrete domain j, i.e., the current height is the previous height plus an input. This structure is analogous to a scalar discrete time dynamic equation. Therefore, discrete domain results will be used to develop a criterion for stability in the layer domain.andDC-Gain Layer-to-Layer Stability CriterionInformally, layer-to-layer stability is essentially analogous to the conventional one- dimensional notion of input-output stability in that it is required that any bounded input result in a bounded output, under some norm.

In the context of (3), the formal definition of layer-to-layer stability is as follows.Definition [5]: The linear repetitive process (3) is said to be layer-to-layer stable if there exists , ensure the real scalars0M0,10jjandwhich, for each constant inputoutput sequence0jjhsatisfies0, 01jjhhMhjwherehis the equilibrium height profile and0his the initial height profile.(3)(4)(5)(6)The definition for layer-to-layer stability requires that the output, in the case of (3)hcontract asymptotically to a layer-independent final value . Physically, Definition 1 requires that the error between the actual height of the part h and the reference part height hr contract asymptotically towards zero each layer.

Physically, Definition 1 requires that the error between the actual height of the part h and the reference part height hr contract asymptotically towards zero each layer. Applying results from both repetitive process control theory and conventional one-dimensional discrete time systems leads to the following theorem.Theorem 1: The DC-gain LMD process in (3) is layer-to-layer stable if211.Examining the definition of the linearization constant κ2 in (5), it is clear that it depends explicitly on the slope of the catchment efficiency function fμ. Therefore, as stated in Theorem 1, the process is locally layer-to-layer stable if the rate of catchment (the slope of the powder catchment efficiency) is positive with higher rates leading to larger stability regions.

Note that while (7) requires both that 0 < κ2 < 2, in practice |κ2| < 1.Powder Catchment Efficiency Influence on Layer-to-Layer StabilityFor a more physical interpretation, consider a structure whose top surface has a notch feature and a powder catchment function which follows the Gaussian shape as illustrated in Figure 3,2,,exp100PjcenterscalePjwidthdfdwhere αscale is the maximum powder catchment efficiency (%), αcenter is the vertical distance below the nozzle at which maximum powder catchment efficiency occurs (mm), αwidth is the powder catchment efficiency characteristic width (mm), and dP,j = dS,j – hj–1 is distance between the part and the cladding head (mm).

In LMD processes, the blown powder ejected by the cladding head is deliberately shaped such that at a focal point an optimal amount of powder reaches the melt pool, resulting in the powder catchment efficiency function fμ.In typical LMD processes, the layer-to-layer change in the substrate standoff distance δdS is a constant amount for each layer based on knowledge of the bead height δh for a given set of process parameters. If the part standoff distance dP,0 initially places the melt pool in Region A (the upper pane in Figure 3), the linearization constant κ2 in (5), which is proportional to the rate of change of powder catchment efficiency with respect to part standoff distance, is positive. That is, the powder catchment efficiency is increasing with respect to increasing part standoff distance. In this case, the depth of the notch feature will reduce because the powder catchment efficiency is inversely proportional to feature height.

In this case, the depth of the notch feature will reduce because the powder catchment efficiency is inversely proportional to feature height. When the substrate standoff distance is incremented a constant amount, the melt pool again lies in Region A, further reducing the depth of the notch feature. If the multi-layer deposition is continued without operator intervention when the melt pool lies in Region A, evidence of the notch will continue to attenuate from layer to layer until a constant part height is achieved, as illustrated by the layer-domain schematic onjh,(7)(8)the right hand side of the top pane in Figure 3. Additionally, as the slope of the powder catchment efficiency function increases, the attenuation of the notch feature occurs in fewer layers.Alternatively, if the part standoff distance dP,0 initially places the melt pool in Region B (the lower pane in Figure 3), the linearization constant κ2 is negative.

That is, the powder catchment efficiency is decreasing with respect to increasing part standoff distance. Incrementing the substrate standoff distance a constant amount and repeating the deposition process with the part standoff distance such that the melt pool lies in Region B further amplifies the notch feature. Without operator intervention, the layer-to-layer growth will eventually lead to catastrophic defects in the deposition, as illustrated in the lower pane in Figure 3.dPhxBValley Receives More Powderfμ ADirection of TravelValley Receives Less PowderPeak Receives More PowderNozzlehxPowder StreamhxLayer12345fμ AdPLayer12345hxPeak Receives Less PowderBFigure 3.

Schematic of deposition process around a stable part standoff distance (Region A, top) and around an unstable part standoff distance (Region B, bottom).From a control theoretic standpoint, setting the part standoff distance dP such that the melt pool lies in Region A is an open-loop layer-to-layer stable operating point. Conversely, a part standoff distance such that the melt pool lies in Region B is a layer-to-layer unstable operating point. Additionally, because the DC-gain linear process in (3) is analogous to a 1D discrete time system, the value |1–κ2| is the DC-gain layer-to-layer pole location. As κ2 approaches unity (higher catchment efficiency slope), the layer-to-layer pole location move towards the origin of the unit disc. The causes the layer-to-layer process speed to increases.

As κ2 approaches unity (higher catchment efficiency slope), the layer-to-layer pole location move towards the origin of the unit disc. The causes the layer-to-layer process speed to increases. This insight follows from discrete-time system theory.Layer-to-Layer Stability Process MapBecause the DC-gain stability criterion in (7) is explicitly dependent on both the spatial flow rate λ and the part standoff distance, a series of process maps can be constructed to aid the selection of process parameters and operating conditions that result in layer-to-layer stable builds. In previous work by the authors [6] the model presented in the Model Description and Preliminaries section was identified for a well-tuned process on an Optomec MR-7 [7] system using 316 Stainless Steel. The resulting model parameters are given in Table 1.Table 1. Open-Loop LMD Process Parameters.

When the stability criterion (7) is violated, the corresponding combination of parameters is deemed an unstable region of the parameter space and is labeled as such. Additionally, because the physical process has limits on how much material can actually be deposited, values of the spatial flow rate which exceed 1×10-3 kg/mm are deemed infeasible. Figure 3 shows the resulting process map around the process operating point given by the parameters in Table 1.In Figure 4, the λ contours represent levels of constant spatial flow rate with zero spatial flow rate occurring in the lower right hand corner of the process map. When spatial flow rate is small, the available values of part standoff and the achievable bead height which yield stable layer-to-layer builds is small, i.e., the length of the contour between dP = 5 mm and dP = 10.5 is shorter for λ = 1×10-5 kg/mm than that of λ = 5×10-5 kg/mm.

As spatial flow rate increases, the region of stability widens significantly until approximately λ = 1×10-4 kg/mm when the stability window begins to decrease as the process approaches the unstable region located on the left hand side of the process map.The second set of contours located on the map indicates constant levels of the stability criterion in (7). The right hand boundary of the stable region indicates the powder catchment center αcenter = 10.57 mm. At this location in the powder catchment efficiency function, there is an inflection point where the slope switches from being positive to being negative. In other words, when the part standoff distance is less than αcenter the process is operating in Region A as shown in Figure 3. When the part standoff distance is larger than αcenter the process is operating in Region B in Figure 3. The left hand boundary is a more process-dependent boundary.

When the part standoff distance is larger than αcenter the process is operating in Region B in Figure 3. The left hand boundary is a more process-dependent boundary. At this location, the desired bead height and spatial flow rate result in a process which builds too fast and would be coincident with the cladding head if no corrective action was taken, i.e., increase the layer-to-layer increment in substrate standoff distance.1080.5604|1-2| Contours1221UnstableRegionUnstableRegion1.5 Contours0.20.20.20.40.40.40.60.60.60.80.80.8111Part Standoff Distance (mm)Bead Height (mm)2.532Figure 4. Layer-to-layer stability process map.Simulation and Experimental ResultsTo demonstrate the efficacy of the stability criterion and the process map discussed in the previous two sections, two simulations are performed using the full nonlinear dynamic model in [4] and corresponding experimental results are presented.

The first simulation and experimental case presented is a process which is operating in Region A with constant process parameters. The second simulation and experimental case is a process which is operating in Region B with constant process parameters.An important use for the LMD process is the repair of high-value parts. In these operations, material may be removed from the part to be repaired in a fashion such that the resulting substrate is non-uniform, i.e., corroded sections are removed while non-corroded sections are left untouched. The result of this operation may result in an uneven substrate onto which the repair is made. In order to simulate this, a portion of the substrate has been machined to leave a pocket 0.600 mm deep by 25.4 mm long. A photograph of substrate and pocket used in the experimental cases is shown in Figure 5.Figure 5.

A photograph of substrate and pocket used in the experimental cases is shown in Figure 5.Figure 5. Photograph of substrate and pocket feature used in experimental cases.The simulations are seeded with this feature as the initial condition for part height. The model parameters are given in Table 1 and the process parameters are listed in Table 2. Figure 6 shows the location of the two operating points, along the λ = 1.8×10-5 kg/mm contour, on the process map. For Case 1, the stable build, κ2 = 4.570×10-2. For Case 2, κ2 = -4.950×10-2 indicating an unstable build. The only difference between the two build cases is the part standoff distance.Table 2.

For Case 1, the stable build, κ2 = 4.570×10-2. For Case 2, κ2 = -4.950×10-2 indicating an unstable build. The only difference between the two build cases is the part standoff distance.Table 2. Process parameters used for cases.Process Parameter Bead Height, δh (mm) Spatial Flow Rate, λ (kg/mm) Part Standoff Distance, dP (mm)Case 1 (Stable) Case 2 (Unstable)0.381 1.878×10 9.8080.381 1.878×10 11.39612Stability Boundaries0.510Case 113|1-2| Contours1UnstableRegionPart Standoff Distance (mm)Bead Height (mm)9111.510Case 228Figure 6. Process map with operating points for Cases 1 and 2 labeled.The top pane in Figure 7 shows the simulated part height for Case 1 at every third layer from layer j = 1 to j = 24.

Process map with operating points for Cases 1 and 2 labeled.The top pane in Figure 7 shows the simulated part height for Case 1 at every third layer from layer j = 1 to j = 24. As evidenced by the height profiles in Figure 7, because the process is layer-to-layer stable, the perturbations from equilibrium caused by the initial pocket in the substrate are attenuated each layer and after approximately layer j = 15, all evidence of the initial pocket is gone and each layer has a uniform bead height. Using the same process parameters, a 316 Stainless Steel thin-walled part was deposited on a steel substrate with a pocket 0.6 mm deep and 25.4 mm long using an Optomec LENS system. A photograph of the resulting deposition is shown in the bottom pane in Figure 7. The qualitative behavior of the deposition and the simulation are in very good agreement.1 Substrate Pocket20050103040Figure 7.

The qualitative behavior of the deposition and the simulation are in very good agreement.1 Substrate Pocket20050103040Figure 7. Simulation height signals for every third layer for j = 1 to j = 24 and photograph of the experimental deposition for stable Case 1.Figure 8 shows the simulated and experimental results for Case 2 where the open-loop process is unstable. While the simulation and experimental depositions start out in a very similar fashion to Case 1 where the open-loop process is stable, the taller features quickly outpace the build rate of the lower features resulting in the large U-shaped defect in the final part. This unstable deposition would either require much earlier operator intervention to move the operating point into the stable region or the taller features would need to be machined off and the deposition restarted.30401001 25.4 mmSubstrate Pocket5020Figure 8.

Simulation height signals for every third layer for j = 1 to j = 24 and photograph of the experimental deposition for stable Case 2.Summary and ConclusionsThe LMD process is an important process for the creation of one-off functional metal components and the repair of high-value parts. In contrast to how it is typically modeled, the LMD process possesses dynamics both within the layer and from layer-to-layer. Because of this, ostensibly valid choices for process parameters can result in failed builds.A two-dimensional repetitive process model of the LMD process is presented which enables the application of repetitive process control theory to develop a DC-gain condition for the stability of LMD fabrications. The criterion is explicitly dependent on the slope of the powder catchment efficiency function. Because the criterion is a function of process inputs, process maps can be generated for specific processes to determine the regions of stability in the process parameters space.

Because the criterion is a function of process inputs, process maps can be generated for specific processes to determine the regions of stability in the process parameters space. The map could enable operators to quickly tune or plan their process builds without the need for extensive testing. Simulation and experimental results of twooperating points – one stable and one unstable – indicate that the process map is effective in predicting the stable and unstable behavior of multi-layer LMD fabrications.AcknowledgementsThis work was supported in part by the National Science Foundation (CMMI1301414) and through a US Department of Education GAANN Fellowship (P200A120062), with technical and experimental support from Optomec and MachMotion.References[1] Doumanidis, C. and Kwak, Y-M., 2001, “Geometry Modeling and Control by Infrared and Laser Sensing in Thermal Manufacturing with Material Deposition,” ASME Journal of Manufacturing Science and Engineering, vol. 123, no. 1, pp.

132, no. 1, pp. 011010:1-9.[4] Sammons, P.M., Bristow, D.A., Landers, R.G., 2014, “Control-Oriented Modeling of Laser Metal Deposition as a Repetitive Process,” American Control Conference, Jun 4-6th, Portland, OR.[5] Rogers, E., Galkowski, K., Owens, D.H., Control Systems Theory and Applications forLinear Repetitive Processes, Berlin, Springer, 2007.[6] Sammons, P.M., Bristow, D.A., Landers, R.G., 2014, “Frequency Domain Identification of a Repetitive Process Control Oriented Model for Laser Metal Deposition Processes,” International Symposium on Flexible Automation, July 14-16th, Awaji Island, Hyogo, Japan.[7] “LENS MR-7 Systems – Optomec Additive Manufacturing,” Optomec AdditiveManufacturing, Web, 10 July 2015..

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA Deep Learning approach to Defect Detection in Additive Manufacturing of Titanium AlloysX. Liu*, A. Mileo*Advanced Manufacturing Research Centre (I-Form), School of Computing, Dublin City University, Dublin, IrelandAbstractIn Additive Manufacturing (AM) of titanium alloys, the formation of defects in parts is typically related to the stability of the melt pool. With increased instability and size of the melt pool comes an increase in the level of emissions generated as the laser processes the material. Recent developments with in-situ monitoring and process control allows the collection of large amounts of data during the printing process. This includes data about emissions, which are made available as 2D representations in the form of colour images.

This includes data about emissions, which are made available as 2D representations in the form of colour images. However, it is still a manual process to inspect these 2D representations to identify defects, which hinders scalability. Given recent advances in Deep Learning for computer vision and the availability of large amounts of data collected from in- situ monitoring, our approach leverages Deep Learning techniques for characterizing abnormal emissions to automatically identify defects during the printing process. One of the challenges to apply deep learning in AM is the lack of proper labelled data for training the models. In this paper, we tackle this challenge by proposing an approach that uses transfer learning and fine-tuning on a pre-trained Convolutional Neural Network (CNN) model called VGG 16 to successfully train the deep model with a small labelled dataset.

Results show good classification accuracy on the emission images obtained from the in-situ monitoring system, and improvements in classification of defects on a public industrial benchmark datasets named DAGM (Deutsche Arbeitsgemeinschaft für Mustererkennung e.V., German chapter of the IAPR).1 IntroductionIn Additive Manufacturing (AM), also known as 3D printing, a laser scans the input material and a melt pool is created at the laser-material interaction point. This forms a layer of the object being manufactured, and the process continues layer by layer until the manufacture is complete. The formation of defects in parts such as tensile weakness, is typically related to the stability of the melt pool during manufacture. Due to thermal instability, the melt pool can create different levels of emissions. A more unstable and volatile melt pool will typically emit a greater number of emissions leading to the formation of defects in parts.

A more unstable and volatile melt pool will typically emit a greater number of emissions leading to the formation of defects in parts. Recent developments in monitoring and process control have resulted in a significant enhancement of the AM process and reduced the amount of inter-build variation and interruption in material manufacturing. Also, given recent advances in computer vision and the availability of potentially large amounts of data collected from in-situ monitoring of emissions from melt pools during additive manufacturing, it is possible to use machine learning on this data to automatically identify key features and predict the presence of defects in manufactured products.In-situ monitoring facilitates the collection of large amounts of data during the build process in AM.

Specifically, this data includes feedback on energy input and emissions from the AM build process, and it is typically done through two sensor modules: one module measures the power of laser input at any point and the other module measures emissions from the melt pool in the near-infrared infrared spectra. The data streams collected from these modules can be used to build 2D and 3D representations of the objects being manufactured. Figure 1 shows one example of the images formed by emissions in 2D and 3D representations produced by the InfiniAM monitoring suite present in the Renishaw machine. Despite the abundance of this type of data, the analysis and characterization of emissions as well as their correlation with defects, is still a manual process that involves examining the representations produced by the monitoring software.

As the performance of humans in such tasks can vary, along with fatigue and the associated costs of labor, there is demand for the automation of such inspection. In addition to the cost of labour, defects in manufacturing usually mean a waste of time, energy,423and other resources. Ideally, a real time prediction of a defect during the manufacturing process could tell the operator to either shut down and stop production at an early stage or intervene, if possible, to prevent further defects from happening. This requires not only to develop models with the ability to better and automatically understand the data and relate it to defects, but it also requires to continuously adjust such models as the data is produced.

This research aims to explore modern data analytics and machine learning techniques including Convolutional Neural Networks (CNNs) to automatically inspect the huge amounts of data captured by an in-situ monitoring suite to extract the most descriptive features that can help in tasks such as classification and object detection. More specifically, in this research we propose a neural network based architecture to accurately and efficiently identify (and classify) input data as normal or abnormal with a limited amount of labelled initial data.The material for the 3D-printed parts considered in this paper is Ti6Al4V. The in-situ monitoring suite is the InfiniAM monitoring suite from the Renishaw 3D printer, which can provide feedback on energy input and emissions using two photodiodes as sensors.

The in-situ monitoring suite is the InfiniAM monitoring suite from the Renishaw 3D printer, which can provide feedback on energy input and emissions using two photodiodes as sensors. The sensors detect plasma emissions (range 700 nm to 1040 nm) and near-infrared (range 1090 nm to 1700 nm) from the melt pool and the monitoring suite produces a 3D point cloud model from which we can look 2D representations, which is a horizontal cut of the 3D model. Using a video recording software, we were able to extract videos of the model representing the building process layer by layer. We extracted each of these layers as one image (440px840p resolution) by removing duplicated frames so that we have one image only per layer in the AM process. These images represent the input for our model and from now on we refer to them as to the emission dataset.

These images represent the input for our model and from now on we refer to them as to the emission dataset. The rest of the paper is organised as follows: Section 2 is a literature review of related approaches to optical inspection in AM using machine learning; Section 3 illustrates the proposed methodology based on CNN and presents our evaluation setup for two experiments; Section 4 discusses the results obtained from the tests and future research directions; we conclude in Section 6 with a summary and final remarks.Figure 1: 3D to 2D representations of emissions during manufacture4242 Literature reviewDuring the AM process, as layers are built up to make a component, the height of each layer and the classification of this height is hugely important as the next layer on top of this will be impacted if height is not within certain tolerances.

Thus, an optical inspection task is required to detect variations in size, orientation of patterns and other characteristics that could reflect any defects in a layer. It is also important to develop flexible inspection approaches that can be easily reconfigured for different tasks. Many of the conventional approaches to optical inspection are based on feature extraction and further clustering or classification. Over the last decade, advances in deep learning have made it possible to apply Neural Network architectures to automatically inspect the surface of the component and the layers printed during the AM process.2.1 Feature based approachesConventional approaches to AM inspection employ human engineered feature extraction to find defective patterns in an image.

The authors of [14] presented a strategy using thresholds derived from a histogram of grey values for segmentation and further computed the shape describing features while in [9] several types of wavelet model have been introduced to extract features from a surface topography. Moreover, [6] proposed an approach for defect detection in texture images with high accuracy by evaluating the distribution of local gradient magnitudes based on a Weibull fit. There are also other methods for feature extraction based on machine learning such as histogram of oriented gradient (HOG) [15], local binary patterns (LBP) [11] and a gray-level co- occurrence matrix (GLCM) [2]. The results generated by these algorithms for feature extraction and defect detection are high in accuracy. However, their algorithms have limitations in their applicability as they often rely on thresholding. Thresholds are sensitive to background, colour and light.

However, their algorithms have limitations in their applicability as they often rely on thresholding. Thresholds are sensitive to background, colour and light. When these conditions change, manual adjustment of the thresholds is needed. Also, such human-engineered methods are limited by the fact that the features need to be customized to specific AM tasks. As a result, for complex AM conditions these methods are not robust and discriminative enough to generate results with sufficient accuracy.2.2 Applying CNNs for inspection in AMIn order to develop approaches that are both good on adaptability and high in accuracy, several methods based on convolutional neural networks (CNNs) have been proposed. [8] and [5] have applied CNNs in their work and achieved higher classification accuracy than conventional machine learning algorithms. However, a major problem is that training a deep CNN from scratch requires a large amount of labelled data.

However, a major problem is that training a deep CNN from scratch requires a large amount of labelled data. One of the challenges we consider in this paper is to be able to apply CNNs to AM processes with acceptable accuracy when only a limited amount of labelled data is available. Authors [1], [13] and [7] have used transfer learning to address this problem by using pre-trained weights from a source network to set the weights of a target network and then use the target network to fulfil the task of feature extraction. However, the performance improvements with these approaches depends on the fact that there is similarity between the source and target domains in their tasks.

However, the performance improvements with these approaches depends on the fact that there is similarity between the source and target domains in their tasks. Authors in [3] have pointed out that if there is a significant difference between the source and target domains such a transfer learning approach with fixed transferred weights can yield less accurate results, and this is the case for AM processes.2.3 Transfer learning with fine-tuningIn order to address the low performance of transfer learning when using dissimilar target and source domains, authors in [10] proposed a method that applies fine-tuning on the CNN architecture VGG 16 [12]. In their approach, both the source and target networks were based on VGG 16. The only modification on the VGG 16 architecture was reducing the 1,000 node outputs (as used in the source network) to 12 nodes on the target network.

The only modification on the VGG 16 architecture was reducing the 1,000 node outputs (as used in the source network) to 12 nodes on the target network. For the data, they used the ImageNet 2012 dataset as the source dataset and the industrial optical inspection dataset DAGM 2007 [16] provided by the German Association for Pattern Recognition, as the target dataset. The results of their experiments showed significant improvements on the overall classification425performance. Building upon this approach, in this paper we develop a methodology to process 2D representations of the emission data from InfiniAM and DAGM for defect detection and analyse changes in performance.3 MethodologyIn this section, the proposed classification method is illustrated and tested using the emission dataset we generated and the DAGM datasets. We first provide a description of the initial dataset and the experimental setup.

We first provide a description of the initial dataset and the experimental setup. Secondly, the performance of two different types of classifiers is investigated. Thirdly, the setup of the deep learning model is modified to improve the classification accuracy and all the datasets are processed by the improved model with a comparison of performance. For reproducibility, we also provide a list of specific parameters and settings used in the overall deep learning model and in the training process.3.1 Initial datasets creationThe first set of data was collected from the 2D representations of the emission images generated by the in-situ monitoring suite. The images represent the melt pool conditions of the printed layers in a group of dog- bone shaped testing parts of Titanium alloy (Ti6Al4V) during the AM process. 11,000 images were manually exported from the InfiniAM monitoring software as described in the introduction. Initially, all the raw image data are unlabelled.

11,000 images were manually exported from the InfiniAM monitoring software as described in the introduction. Initially, all the raw image data are unlabelled. These raw data are not suitable to be used directly for training. For initial inspection, the 11,000 sample images were manually inspected in order to select 150 images as defected samples and another 150 images as normal samples. These were labeled and used to create a dataset considered as the ground truth in the tests. The size of this labelled dataset is relatively small for the training of the ML models. We acknowledge that the manual selection of the samples could be better analysed and automatic methods could be used to select the most representative sample or investigate the impact of sample selection in the training process.

However, in this investigation, we aim at demonstrating the feasibility of the approach with limited training data and we postpone the analysis of sample selection to future work.As the currently available labelled data from the emission dataset are limited, we also used a well-known industrial optical inspection dataset provided by DAGM for testing purposes. The DAGM dataset contains 6 patterns of texture with each texture pattern containing 1,000 non-defective and 150 defective images, resulting in 6,900 images of 12 classes. Based on the original DAGM dataset, from each pattern, 150 non-defective images samples were selected from the original 1000 non-defective samples to create new groups of testing dataset that involves 6 patterns and each pattern include 150 defective and 150 non-defective samples. This dataset was used in conjunction with the manually labeled emission image dataset generated from the InfiniAM 3D model.

This dataset was used in conjunction with the manually labeled emission image dataset generated from the InfiniAM 3D model. In total, there are 7 patterns used in the experiments: along with the 6 patterns from the DAGM dataset, the emission dataset is used as the 7th pattern and it contains samples of both defected and non-defected classes. Figure 2 shows a collection of examples for normal and defect from each pattern. All the tests in the following sections are based on these datasets.426Figure 2: Sample images marked as pattern 1-6 from the DAGM industrial optical inspection dataset.The 6 different texture patterns contain both normal and defects in a particular location, marked with red ellipsoids in the figures.

Images marked as pattern 7 are those we generated from the InfiniAM emission dataset.3.2 Combining VGG 16 and SVM classifierIn this test, the image datasets are used as input for features extraction, The features are extracted by the CNN architecture from a VGG16 model using transferred weights trained using ImageNet data [4] without fine- tuning. Then, features are passed to a Support Vector Machine (SVM) classifier to detect if the input image is normal or has a defect. The architecture of this model is shown in Figure 3.Figure 3: The architecture of the model using a combination of VGG 16 and SVM.427The SVM classifier divides the dataset into classes by creating hyperplanes in multidimensional space. The SVM classifier in this paper is implemented using the C-Support Vector Classification functions from the Python ML package Scikit-learn [18].

C is the penalty parameter, which is used to maintain regularization. It represents misclassification or error term of the SVM classifier. The kernel coefficient for the RBF, the value of Gamma depends on the number of features from the dataThe 7 patterns were individually passed through the model for training and testing purposes. As illustrated previously, the dataset includes 300 samples for each pattern, which are further divided into 2 classes: 150 positive samples and 150 negative samples. 70% of the samples were used for training and the remaining 30% were used for validation.

70% of the samples were used for training and the remaining 30% were used for validation. The classification results are shown in Table II.Table II: classification results from the SVM classifierClassPrecisionRecallF1-scoreSupportClass1 Def0.880.150.2542Class1 Class2 Def Class2 Class3 Def Class3 Class4 Def Class4 Class5 Def Class5 Class6 Def Class6 Class7 Def Class70.50 0.50 0.57 0.47 0.00 0.47 0.00 1.00 0.84 0.57 0.94 0.91 0.980.98 0.52 0.54 1.00 0.00 1.00 0.00 0.79 1.00 0.92 0.35 0.98 0.920.66 0.51 0.55 0.64 0.00 0.64 0.00 0.88 0.91 0.72 0.52 0.94 0.9548 42 48 42 48 42 48 42 48 42 48 42 48As the size of each labelled dataset is quite small, it is remarkable to see how, even with limited training data, the results of classification for certain classes (such as class 7) have relatively high value in recall with true positive rate (precision) and true negative rate (recall) being 98% and 92% respectively.

Despite these positive results, the approach still shows low performance on classes related to patterns 1,2,3 and 4. There are three possible reasons to explain such a low performance: first, the number of samples for training is not sufficient to train the SVM classifier; second, some of the features extracted from the datasets may not be representative enough and may contain too much unnecessary information; third, SVM may not be a suitable classifier for certain patterns in the test data.

To address these issues and improve results, we propose and test a different approach solely based on CNN, which relies on VGG 16 architecture extended and modified for our specific problem.4283.3 Extending VGG 16 for classification using dense layersThis approach relies on a transfer learning method in which the 13 convolutional layers from the pre- trained convolutional layers of VGG16 from the previous model are still used for feature extraction and the weights in these layers are unchanged. To modify the classifier, the SVM classifier is replaced by fully connected layers.

To modify the classifier, the SVM classifier is replaced by fully connected layers. As illustrated in Figure 4, after the convolutional layers, 2 dense layers with ReLU activation function are added and followed by 1 dense layer as the output layer using sigmoid as the activation function, since detecting normal or defect individually for each pattern isa binary classification task.Figure 4: The architecture of the model based on VGG 16 and fully connected layers.In this experiment, rather than directly test on all the image groups, three patterns which are considered to be the most challenging are selected first.

As illustrated in Table II, the selected patterns are: pattern 4, with one of the worst performances in the previous test; pattern 6, with average performance in the previous test; pattern 7, with the best performance in the previous test.Test on pattern 6: The initial investigation of the performance of the new model began with the dataset of pattern 6 (refer to Figure 2) for binary classification and the length of training in this test was set to 200 epochs. Accuracy and loss are the metrics used to trace and evaluate the training and validation process. Figure 5 shows the curves of accuracy and loss for both training and validation for pattern 6.In Figure 5, the lines show an overall increase in accuracy (the left image) and a decrease in loss (the right image) both for training and validation but with significant fluctuation. The results do not show overfitting as the validation loss is not significantly larger than the training loss.

The results do not show overfitting as the validation loss is not significantly larger than the training loss. Regarding the problem of the extremely high fluctuation in the validation, the following points are considered:1) The number of samples for the test is relatively small, as the loss is still relatively high and unstable, even a small number of classification results will cause major changes in the overall accuracy.2) There may be too much noise considered as features. To address this problem, adding a pooling layer between the convolutional and the dense layers would be a solution.3) According to Kim et al. (2017) [10] with a frozen network for the convolutional layers, the results should have a level of accuracy in the range of 78%-85%. This means that using the setup without fine-tuning, the fluctuation should be around 80%.

This means that using the setup without fine-tuning, the fluctuation should be around 80%. Fine-tuning has therefore been applied as discussed in Section 3.4 to reduce the fluctuation and improve the overall classification accuracy.429Figure 5: Accuracy and loss over 200 epochs (the horizontal axis shows the number of epochs) with transfer- learning and without fine-tuning. The vertical axis in the left chart stands for the value of accuracy while the vertical axis in the right chart stands for the value of loss.3.4 Further optimisation: pooling and fine-tuningTo further improve the model, an average pooling layer is added between the convolutional layers and the dense layers to reduce the spatial dimension of the output from the convolutional layers.

We performed a test by training the new model for only 50 epochs to show that the modified architecture can effectively and quickly reduce the fluctuation for accuracy and loss in both training and validation, as shown in Figure 6.Figure 6: There is an obvious reduction of fluctuation in the curve of the overall accuracy and loss after adding the pooling layers (top-right and bottom-right) compared to the results without pooling a layer (top-left and bottom left).After improving the model with the addition of a pooling layer, fine-tuning is applied in order to increase the classification accuracy and reduce the loss. Based on this new configuration, the last 3 convolutional layers from the VGG16 architecture are unlocked for fine-tuning. The weights in the unlocked layers are modified by the training replications using the image datasets. Using the updated model and pattern 6, the training and validation are re-executed for 200 epochs to ensure the process reaches convergence.

Using the updated model and pattern 6, the training and validation are re-executed for 200 epochs to ensure the process reaches convergence. The curves of training and validation as shown in Figure 7. Around epoch 200, the training accuracy is over 97% while the validation430accuracy is around 95%. The training loss starts to converge at about the 100th epoch with a value around 0.1 while the training loss continues to drop slightly.Figure 7: Accuracy and loss for pattern 6, using a model with pooling and fine tuning on the last 3convolutional layers.Test on pattern 4: Similar tests are also applied on pattern 4, with the updated setup, the validation accuracy at about the 200th epoch is also over 96% with a relatively stable value of the validation loss of 0.1 (refer to Figure 8).

This result is significantly more accurate than the one obtained from the SVM classifier in the previous experiment, where the model was not able to identify the classes for pattern 4.Figure 8: Accuracy and loss for pattern 4, using a model with pooling and fine tuning on the last 3 convolutional layers.Test on pattern 7: Pattern 7 is the emission dataset of images generated from the 2D presentation collected and converted from the InfiniAM in-situ monitoring system. This dataset is the one we specifically collected for this research. The model is trained using the dataset for pattern 7 for 200 epochs. The output of the classification shows considerably high values in accuracy for both training and validation.

The model is trained using the dataset for pattern 7 for 200 epochs. The output of the classification shows considerably high values in accuracy for both training and validation. As shown in Figure 9, the training process reaches convergence after around the 130th epoch with a training accuracy around 99% and a validation accuracy over 98%.431Figure 9: Accuracy and loss for pattern 7, using the model with pooling and fine tuning on the last 3 convolutional layers.The results of the classification by the model with fine-tuning has shown significant improvement in the overall accuracy. To complete this investigation, the remaining patterns 1, 2, 3 and 5 (refer to Figure 2) are also tested individually using our modified VGG16 architecture with transfer learning and fine-tuning.

To complete this investigation, the remaining patterns 1, 2, 3 and 5 (refer to Figure 2) are also tested individually using our modified VGG16 architecture with transfer learning and fine-tuning. Results for all the 7 patterns are shown in Table III, including the average values (after convergence) of training accuracy, training loss, validation accuracy, validation loss and the training duration (in epochs) for each corresponding pattern.Table III: classification results from the modified VGG16 architecture with transfer learning and fine-tuningPatternsPattern 1 Pattern 2 Pattern 3 Pattern 4 Pattern 5 Pattern 6 Pattern 7Ave. Training Accuracy0.98 0.99 0.98 0.97 0.99 0.97 0.99Ave. Training Loss0.03 0.02 0.04 0.05 0.02 0.05 0.03Ave. Validation Accuracy0.96 0.98 0.96 0.96 0.98 0.95 0.98Ave.

Figure 10 describes the architecture of the model layer by layer, and the shape of input/output tensors in each layer. Table IV contains the hyperparameters used for the training of the model.Stopping criterion: In the tests, the numbers of epochs for stopping criterion are simply set as long enough for the curves of validation accuracy and loss to reach the convergence where the curves will not have significant change with longer duration of training. Because the input dataset is different in each test, the number of epochs432required will be different.

For example, to reach the convergence in the training process, the test of dataset 1 requires more epochs than that required by the test of dataset 7.Table IV: hyperparameters used for the training of the deep learning modelNameType / ValueDescriptionOptimizerLoss functionLearning rateEvaluation metricStochastic Gradient Descent (SGD)binary cross entropy0.001Accuracy, LossOptimizers are used to change the attributes of the neural network to reduce the losses The loss function computes the quantity that a model should seek to minimize during training The step size at each iteration while moving toward a minimum of a loss function during the training process Metric is a function to judge the performance of the modelPooling layer: To reduce the spatial dimension of the output of the feature extraction model based on VGG 16, a Global Average Pooling 2D (GAP2D) layer is added between the outcome of the VGG 16 model used for feature extraction and the dense layers of the classifier.

The GAP2D layer receives the output tenser from the VGG 16 model and applies global average pooling operation for spatial data. By doing this the number of total channels in the output of the GAP2D layer is reduced to 512 and ready to be processed in the dense layers. (See Figure 10)Dense layers: We added three dense layers (also known as fully connected layers) to our model. The activation function used in the first two layers is the ReLU activation function, while a Sigmoid activation function is used in the third dense layer for binary classification.Figure 10: The overall model including the VGG 16 architecture layers, the GAP2D layer and the dense layers for classification.

The value “None” in the shape of tensors indicates the dimension is a variable4334 DiscussionFrom the results obtained, it is clear that the deep learning model introduced in our approach which leverages transfer learning and fine-tuning has relatively good classification performance on all the 7 patterns in the initial datasets. For certain patterns, such as pattern 5 and pattern 7 the results show higher values in accuracy even before any optimisation. More importantly, by replacing the SVM with fully connected layers as a classifier, the approach allows the overall model to be fine-tuned using a relatively small amount of data to further adjust the weights in both the convolutional and fully connected layers for low performance patterns.

The application of fine-tuning greatly improved the accuracy in the classifications for those patterns and it is also expected to allow the model to be continuously improved as the available datasets grow in size.With the aid of transfer learning, this approach greatly reduces the required amount of labelled data and computational power for training. In fact, manually creating large training databases is time consuming, expensive, and often infeasible in industrial production settings. The lack of properly labelled data is a common issue for the application of ML, especially in AM processes.Future research we are planning to investigate goes into two main directions. The first direction is looking to continue to improve the deep learning model. Although the current model works well with the initial datasets, the architecture used in this paper is not the most up to date CNN.

Although the current model works well with the initial datasets, the architecture used in this paper is not the most up to date CNN. Other architectures, such as ResNet [17], have better performance when the problem becomes more complex and the network needs to be deeper. This will be considered for the application of computer vision with transfer learning and fine-tuning on other tasks. The second direction for investigation considers the need to produce more high quality data and we are looking into extending our framework to include active learning and semi-supervised learning for automatic data labelling.5 ConclusionIn this paper we have designed and tested a deep learning model based on Convolutional Neural Networks to automatically identify defects in the AM process of titanium alloy Ti6A14V. In our approach, we leverage information about emissions collected through in-situ monitoring and represented as images, one for each layer.

In our approach, we leverage information about emissions collected through in-situ monitoring and represented as images, one for each layer. These images represent a dataset we have curated and used to train our deep learning model. Our experiments have demonstrated that our model pre-trained and fine-tuned can obtain good performances even with relatively low computational power and limited training data. The model we propose in this paper can be used as an effective feature extractor and classifier with limited labelled data for training. However, we believe the ability to generate more labelled data using the outcome of our model is necessary, as it not only enables faster convergence (with limited number of epochs), but it also represents a valuable resource to be used by other researchers.

For this reason, we are investigating the combination of our model with active learning techniques to develop a framework that can produce good quality labelled data to be used when applying machine learning to AM processes. As a result, we will produce a more accurate and more effective pipeline that can potentially be used for different classification tasks where the availability of labelled dataset is scarce. Currently we are looking into two of such tasks: automatic segmentation of individual melt pool areas and characterisation of porosity structure in AM processes.Acknowledgements:This research is supported by a research grant from Science Foundation Ireland (SFI) under Grant Number 16/RC/3872 and is co-funded under the European Regional Development Fund and by I-Form industry partners.434REFERENCES[1] Matthew D Zeiler and Rob Fergus. Visualizing and understanding convolutional networks. In European conference on computer vision, pages 818–833. Springer, 2014.

How transferable are features inDaniel Weimer, Bernd Scholz-Reiter, and Moshe Shpitalni. Design of deep convolutionalFabian Timm and Erhardt Barth. Non-parametric texture defect detection using weibullXiang Jiang, P Scott, and D Whitehouse. Wavelets and their applications for surface435.

fect formation in EBM parts built in horizontal orientationKhalid Rafi. H, Karthik N.V, Thomas L. Starr*, Brent E. StuckerDepartment of Industrial Engineering, *Department of Chemical Engineering, J. B. Speed School of Engineering, University of Louisville, Louisville, KY 40292AbstractEBM built parts in a horizontal orientation, particularly with bottom curved surfaces, are prone to variations in dimensional accuracy and defect formation. These defects likely occur due to the nature of the supports utilized and the build strategy. To understand the effect of support structures and scan direction on these defects, a series of cylindrical parts were built using different diameters, support structures and scan strategies. The as-built samples were mechanically tested and sample cross sections were analyzed. Pore formation and balling effects were observed in the lower section of some samples.

The as-built samples were mechanically tested and sample cross sections were analyzed. Pore formation and balling effects were observed in the lower section of some samples. The study looks at the effects of supports, geometry and scan strategy on the minimization of these defects and improving the dimensional accuracy of horizontally built samplesIntroductionAdditive manufacturing techniques offers unlimited design flexibility to produce almost any geometry [1]. Complex internal features with graded material compositions can be fabricated without much difficulty [2]. Support structures are created based on the part position and geometry. These support structures provides geometrical conformity and structural integrity to the actual part. Support structures are essential when it comes to powder- bed based metal additive manufacturing processes such as selective laser melting (SLM) and electron beam melting (EBM).

Support structures are essential when it comes to powder- bed based metal additive manufacturing processes such as selective laser melting (SLM) and electron beam melting (EBM). The material mass and the surface tension of the molten metal produced by high energy beams could lead to non-homogenous material consolidation at regions of varying cross sections. This could be more prominent when the parts are built in the horizontal orientation where the build orientation and beam scan directions are in the same plane. While building these dimensional complex structures there are flat/curved surfaces facing the horizontal build plane. These surfaces/curvatures are built by providing support structures. Although the powder-bed supports the part to a certain extent, it is not adequate to maintain the structural integrity.

These surfaces/curvatures are built by providing support structures. Although the powder-bed supports the part to a certain extent, it is not adequate to maintain the structural integrity. The support structures in EBM serve two main purposes; to maintain dimensional stability and to provide better conductivity for the electrons from the part to the build plate.Horizontally built samples offer two main advantages. The number of layers required for building them is lower and they offer relatively better strength compared to vertically built samples [3].While doing some preliminary experiments cylindrical specimens in EBM have been fabricated and it was found that the specimens built in horizontal orientation have some differences in final geometry. Examination of the specimen cross-sections also revealed the presence of internal defects close to the bottom surface.

Examination of the specimen cross-sections also revealed the presence of internal defects close to the bottom surface. This became the motivation for a detailed study on factors which cause inaccuracy in the final geometry and the formation ofdefects. In this study cylindrical parts were fabricated in Arcam EBM machine with different diameters, support structures, layer thicknesses and at X, Y planes in the build plate in horizontal orientations. Detailed descriptions about the working principles of EBM system are given elsewhere [4]. Measurements were carried out to find the variance in actual diameter from the design diameter. Microstructural characterization was carried out to analyze the kind of internal defects formed.In this study the process parameters namely the beam current, velocity, focus of the beam, type of raster, base temperature, build chamber pressure, etc. were not taken into account.

were not taken into account. It must be noted that these process parameters greatly affect the build quality. The purpose of this study is to understand how defects can be reduced quickly by just changing the way the parts are set up and designed. Often times it is very hard to understand the exact nature of the problems in the ARCAM machine and we discuss what possible quick solutions one can implement to get more acceptable parts.Materials and MethodsTi-6Al-4V powders with average diameter of 60 µm were used to fabricate the parts. The powder was supplied by Arcam AB(Sweden) and the parts were fabricated in a Arcam S400 EBM machine. The preset process parameter theme in Arcam machine for Ti-6Al-4V material was used for fabricating the parts, with changes in processing conditions. Cylindrical parts of diameters of 4 mm, 8 mm, 12 mm and 16 mm and length of 40 mm were fabricated in horizontal orientations of X and Y axis directions.

Cylindrical parts of diameters of 4 mm, 8 mm, 12 mm and 16 mm and length of 40 mm were fabricated in horizontal orientations of X and Y axis directions. The parts were fabricated in two batches, where the first batch contained parts with 70 µm layer thickness and the second batch contained parts with 50 µm layer thickness. The experimental matrix is shown in Table 1. Corresponding arrangement of specimens in build plate with support structures generated by Materialise Magics software is shown in Fig.1. Other than straight cylindrical parts, tensile and fatigue samples were also built to standard dimensions and the fracture surface after testing was analyzed for defects.Support structures in the Magics software have several preset values corresponding to different kinds of supports known as point, line, area/box and volume supports. For getting parts without curling it is essential to have at least line supports.

For getting parts without curling it is essential to have at least line supports. Having volume supports (Sintering the powder around the part to make it act as a rigid support) has been proven to be bad for post processing of the part since to provide enough support the powder around the part needs to be excessively sintered. ARCAM advises using either the line or the box supports while building the parts, depending on the intricacies present in the part. Often times it was just common practice to use the default line supports.Line supports are characterized by a line along the larger dimension of the part and ribsperpendicular to this central line. The number of ribs is 5 if the length of the central line is < 200 mm and is equal to 12 if length >200 mm.

The number of ribs is 5 if the length of the central line is < 200 mm and is equal to 12 if length >200 mm. Number of teeth on the ribs is 5 if the width of the part is <200 mm and 12 if the width of the part is >200 mm.Box supports are characterized by a box along the edge of the part and several ribs at an angle of 30o going from one edge of the box to the other. The number of teeth along the larger and the smaller edges of the box are either 5 or 12 depending on if the dimension is less than or greater than 200 mm.Table 1. Experimental matrix for Ti-6Al-4V EBM parts built in horizontal orientation70 µm50 µmLine supportBox supportLine supportBox supportX direction Y direction X direction Y direction X direction Y direction X direction Y direction4 mm, 8mm, 12mm, 16mmThe dimensions of the as-built parts were measured and correlated with different parameter sets.

A statistical analysis was carried out to find the effect of processing conditions on the dimensional accuracy. Analysis of Variance (ANOVA) and the corresponding plots were made using Minitab software ver.16. All the fabricated parts were sectioned and the cross- sections were examined using an optical microscope. The specimens were prepared following standard metallographic specimen preparation methods.a)b)Fig.1. Specimens arranged in the build plate with support structures generated by Magics software. a) Dense box supports b) Thin line supportsResults and DiscussionsDimensional accuracyThe parts built in horizontal orientation using different processing conditions wereanalyzed for dimensional accuracy with respect to the actual design diameter. The cross- sectional view of the cylindrical rods resembled a ‘tear drop’ as shown in the schematic (Fig.2).

The cross- sectional view of the cylindrical rods resembled a ‘tear drop’ as shown in the schematic (Fig.2). Maximum and minimum offset in diameters of the ‘tear drop’ structure were measured to find the deviation from the design diameter. The variation from actual diameter or error is plotted against the four main factors namely the type of support, build layer thickness, size of the sample and orientation of the sample. The plot showing the main effects in Fig. 3 and Fig. 4 illustrate how each of the factors affects the percentage error. Lower is the value of the response better is the dimensional accuracy achieved.Fig 2. Schematic showing the tear drop shaped cross-sectional appearance of horizontal built cylindrical sample by EBM on the left and tear drop shaped samples on the rightThe ANOVA result for maximum offset diameter is shown in Table 1 and the corresponding plot is shown in Fig 3.

The P value from ANOVA results indicate that at 95% confidence interval the effect of type of support, build layer thickness and the size of the sample are statistically significant while the orientation of the sample does not hold much significance with respect to the maximum offset diameter.Table 1. Analysis of Variance for Max Offset ResponseSource DF Seq SS Adj SS Adj MS F P Support 1 8.422 8.422 8.422 5.42 0.028 Build 1 21.057 21.057 21.057 13.55 0.001 Size 3 178.488 178.488 59.496 38.28 0.000 Orientation 1 2.280 2.280 2.280 1.47 0.237 Error 25 38.852 38.852 1.554 Total 31 249.100Main Effects Plot for Max Offset Response Data MeansSupportBuild864n a e M21Size21286421234Fig.3.

Main effects plot for maximum diameter offset (Support 1-line support, support 2-Box support; Build 1- 70 micron layer thickness, Build 2- 50 micron layer thickness)From the analysis it can be deduced that by using the following processing conditions the ‘maximum diameter offset’ can be minimized: Support of type 2 (Box supports), Build layer thickness type 2 (50 micron layer thickness), a sample size of type 1 (16 mm diameter or the largest diameter).The ANOVA results for minimum offset are shown in Table 2 and the corresponding plot is shown in Fig 4. The ANOVA results indicate that the support and build orientation have no significance while the build layer thickness and the size of the parts play a significant role. At 95% confidence interval only the build layer thickness affects the minimum offset diameter.Table 2.

Although cross-sectional analysis was carried out for all the 32 samples fabricated, results are given only for those samples which showed maximum/minimum defects. Fig.5 shows some typical defects observed in the cross-section of specimens fabricated in horizontal orientation. An as-built specimen of 4 mm diameter fabricated with line support and 70 µm layer thicknesses shows significant defects (Fig.5a). A specimen fabricated in the same lot and machined to obtain circular symmetry still shows a large defect near to the periphery (Fig. 5b). Relatively better cross-sectional surface can be observed for specimen of 4 mm diameter fabricated with box support and layer thickness of 50 µm (Fig.

5b). Relatively better cross-sectional surface can be observed for specimen of 4 mm diameter fabricated with box support and layer thickness of 50 µm (Fig. 5c).( a )( b )( c )Fig.5 Optical micrograph of as-polished cross-sections of 4 mm diameter a) with linesupport-as built b) with line support-machined c) with box support-machinedThe fracture surface of tensile and fatigue samples with reduced cross-sections of diameter 4 mm also showed severe defects concentrated towards the bottom surface. The fracture surfaces are characterized with deep cleavages containing partly melted powders sticking to the walls. Fig.6 shows the SEM secondary electron image of a tensile fracture surface of an as-built EBM horizontal built sample. The large defects present in the sample can facilitate premature failure during tensile testing. Similar defect patterns can also be observed in the fracture surface of a horizontally built and machined fatigue sample (Fig.7).

Similar defect patterns can also be observed in the fracture surface of a horizontally built and machined fatigue sample (Fig.7). In the fatigue sample, in addition to the cleavage type defects, defects separating some layers are also visible. The nature of the defects indicates that it is an outcome of ‘balling’. Melt ball formation occurs due to melt pool instabilities. This condition is affected by an inadequate energy density transmitted from the electron beam into the powder and results in the so-called ‘balling effect’. Thereby, surface tension of the molten liquid exceeds the wetting ability of the previously solidified layer [5]. This melt balls prevent the continuity of the process and cause the formation of voids which run in length across the sample.

This melt balls prevent the continuity of the process and cause the formation of voids which run in length across the sample. Compared to vertically built samples this condition is much more severe in the horizontal built samples because of the longer beam scans.Further analysis of samples fabricated with larger diameters, box type support and 50 µm layer thicknesses showed cross-sections with very minimal defects. Fig. 8a shows an optical micrograph of cross-section of a cylindrical bar with 12 mm diameter. Compared to the defects observed in lower diameter bars, the defects observed in this specimen are quite negligible. The cross- section of a 16 mm diameter bar (Fig.

Compared to the defects observed in lower diameter bars, the defects observed in this specimen are quite negligible. The cross- section of a 16 mm diameter bar (Fig. 8b) is found to be free from any defects which implies that as the diameter of the curvature increases the possibility of defect formation is minimized.Fig.6 SEM image showing defects closer to the lower curvature at the reduced section in a fractured tensile sampleFig.7 SEM image showing balling defects closer to the lower curvature at the reduced section in a fractured fatigue sample( a )( b )Fig.8.

Optical image at lower part of the cross-sections a) 12 mm diameter bar built with box support b) 16 mm diameter bar built with box supportFrom the dimensional accuracy and defect analysis results it is understood that the nature of the support structure, diameter of the specimen (radius of curvature for specimens other than cylindrical), and the layer thicknesses have significant role in obtaining a horizontal built part with good geometric tolerance and parts free from defects. The dimensional accuracy and the defects were found to be more in lower diameter parts irrespective of the processing conditions. This can be explained based on the mechanism of formation of layers both in the smaller diameter and the larger diameter specimens. Fig. 9 shows the schematic of a smaller diameter and a larger diameter cylindrical cross-section. In a smaller diameter specimen the width of the first few layers is less such that while cooling down, the material tends to shrink towards the center.

In a smaller diameter specimen the width of the first few layers is less such that while cooling down, the material tends to shrink towards the center. The same shrinkage tendency will be there for larger diameter specimens as well. But due to the larger layer width the effect of shrinkage will not be significant. Thus for a smaller diameter it will take more layers for stabilizing the effect of shrinkage when compared to the larger diameter as shown in Fig.9a and Fig.9b. This implies that in the absence of proper supports, the shrinkage due to cooling causes the first few layers to preferentially form a tear drop structure. This will be more prominent in smaller diameter parts. The defects observed in the cross-section can be attributed to an outcome of shrinkage which leads to balling phenomena.Fig. 9.

This will be more prominent in smaller diameter parts. The defects observed in the cross-section can be attributed to an outcome of shrinkage which leads to balling phenomena.Fig. 9. Comparison of layer build-up between smaller diameter and larger diameter cylindrical barsTherefore, to maintain dimensional accuracy it is required to provide dense support structures covering the surfaces up to half the diameter for smaller diameter parts. Whereas, for larger diameter parts dense support structures are required only to support first few layers and the remaining curved surfaces can be supported by lean support structures as shown schematically in Fig 8c and Fig. 8d. Green colored layer in Fig. 8a and Fig. 8b indicates that from that layer onwards there is enough material to support subsequent layers without loss of curvature.Statistical Significance of resultsThe ARCAM machine uses an alternating scan strategy where each layer the direction of scanning changes by 90o.

This means that the X and the Y orientation practically give us the same results, which is evident from the p- values of the dimensional accuracy. Thus for all studies it can be considered that there were two replicates. In several samples two sections were observed and it has been found that if there is a defect in one section often times it is found in the other sections of the same sample as well. As far as the dimensional accuracy goes it is well measurable and the statistical significance can be clearly seen from the results. In the case of defect reduction it is a qualitative conclusion that we observed better mechanical properties [3] and far less number of defects in all the observed sections with the specified build conditions.ConclusionsFollowing conclusions can be drawn from the current study,1. Better dimensional accuracy can be maintained in horizontal built samples by using box support structures and lower layer thicknesses.2.

Better dimensional accuracy can be maintained in horizontal built samples by using box support structures and lower layer thicknesses.2. The effect of orientation of samples in X or Y directions in the build plate is negligible. 3. The defect formation in horizontally built samples can be attributed to the shrinkage issues and the ‘Balling’ phenomena.4. The defects can be minimized in a horizontal built sample by using box type support structures and 50 µm layer thickness.5. The tendency of defect formation decreases with increase in the radius of curvature.AcknowledgementsThe authors acknowledge ONR for support through grant #’s N00014-09-1-0147, N00014‐10‐1‐0800, and N00014-11-1-0689.References1. M. F. Zha, S.Lutzmann. Modeling and simulation of electron beam melting. Prod. Eng. Res. Devel. (2010) 4, pp.15–232. L.E. Murr, E.V. Esquivel, S.A. Quinones, S.M. Gaytan, M.I. Lopez, E.Y. Martinez, F. Medina, D.H. Hernandez, E. Martinez, J.L. Martinez, S.W.Stafford, D.K.

fect Morphology in Ti-6Al-4V Parts Fabricated by Selective Laser Melting and Electron Beam MeltingHaijun Gong*, Khalid Rafi*, N.V. Karthik *, Thomas Starr†, Brent Stucker*Department of Industrial Engineering, †Department of Chemical Engineering, J. B. Speed School of Engineering, University of Louisville, Louisville, KY 40292 Accepted August 16th 2013AbstractIn order to investigate the morphology of defects present in Selective Laser Melting (SLM) and Electron Beam Melting (EBM) processes, Ti-6Al-4V specimens were fabricated with varying porosity using non-optimum processing parameters. Defective specimens were sectioned and polished for microscopy. Image processing was adopted for statistically analyzing the characteristics of defects, such as distribution of defect area and dimensional proportion of each defect. It is found that defect morphology is influenced by process parameters as a result of a variation in the melt pool.

It is found that defect morphology is influenced by process parameters as a result of a variation in the melt pool. Image processing of a cross-section could be a feasible way for calculating porosity of specimens.1. IntroductionSelective Laser Melting (SLM) and Electron Beam Melting (EBM) are two common powder bed fusion processes within Additive Manufacturing (AM) for fabricating parts from metallic powders [1]. The SLM process selectively melts metallic powder using a focused laser beam. The melted powder attaches to the previous layer or support structure and solidifies in a short time. After that, a layer of new powder is recoated upon the current layer for the following layer melting [2]. At present, multiple materials can be utilized for SLM such as stainless steel, maraging steel, cobalt chromium and titanium alloys. The EBM process is similar. But, an electron beam, which is generated by a tungsten filament in the electron beam gun, is used instead of a laser beam.

The EBM process is similar. But, an electron beam, which is generated by a tungsten filament in the electron beam gun, is used instead of a laser beam. When the electrons penetrate the powder surface and further into the powder grains, their kinetic energy is converted into thermal energy [3]. Titanium Ti-6Al-4V and cobalt chromium powder are well developed for EBM.Numerous studies on SLM and EBM have been carried out for process and material development. Additively manufactured materials were also tested and compared to that made from conventional methods. Due to their distinguished layer adding process and rapid cooling rate, SLM and EBM materials exhibit special microstructures and outstanding mechanical properties. Ti-6Al- 4V powders are widely used in SLM and EBM processes for their inherent properties of fracture resistance, fatigue behavior, corrosion resistance and biocompatibility [4].

Ti-6Al- 4V powders are widely used in SLM and EBM processes for their inherent properties of fracture resistance, fatigue behavior, corrosion resistance and biocompatibility [4]. However, defects can be easily formed in additively manufactured Ti-6Al-4V material depending upon the influence of multiple factors. Thus, an investigation of defect morphology is helpful for understanding their effects on material properties.AM powder bed fusion processes can be regarded as a type of Powder Metallurgy (PM) process [5], which are commonly used to produce porous parts. When an AM process is utilized to fabricate porous parts, it can be used to create structures which exhibit extreme flexibility in structural diversity and porosity characteristics if the morphology and distribution of defects in Ti- 6Al-4V material is well-understood.This paper studies the dimensional distribution, morphology of defects, and porosity generated by varying processing parameters in SLM and EBM.

The reasons for defect formation are discussed. Variations in defect morphology with different energy input are analyzed. The statistical results are thus a good reference for future comparisons with traditional PM-produced porous materials.2. Material and method2.1 Powder PropertyRaymor Ti-6Al-4V powder (Grade 23) with an apparent density of 2.55 g/cm3 was used for SLM process. It has a particle size distribution between 17.36 µm (D10) and 44.31 µm (D90) with Mean Volume Diameter around 30 µm. Lots of fine particles are included in Raymor Ti-6Al- 4V powder. Arcam Titanium Ti-6Al-4V ELI powder was used for EBM process. Its apparent density is no more than 2.7 g/cm3. The particle size is normally distributed between 46.94 µm (D10) and 99.17 µm (D90) with a Mean Volume Diameter around 72.69 µm.

Its apparent density is no more than 2.7 g/cm3. The particle size is normally distributed between 46.94 µm (D10) and 99.17 µm (D90) with a Mean Volume Diameter around 72.69 µm. Both Raymor and Arcam powder particles are spherical.2.2 Fabrication of Ti-6Al-4V SpecimensPorosity (RD) of additively manufactured Ti-6Al-4V specimens can be estimated by measuring density using Archimedes method, and then compared with nominal density of Ti-6Al- 4V material. Lower measured densities result in larger porosities.A process window, as shown in Fig. 1, illustrates the effect of various laser power and scanning speeds when using SLM to process Raymor Ti-6Al-4V powder. Processing parameters can be divided into four melting zones. Zone I parameters result in nominally fully dense specimens. Specimens built at a high energy densities with Zone II parameters result in over melting. Defects are often generated in these specimens.

Zone I parameters result in nominally fully dense specimens. Specimens built at a high energy densities with Zone II parameters result in over melting. Defects are often generated in these specimens. Ti-6Al-4V powder is not completely melted using lower energy Zone III parameters, which results in pores and voids. Specimens cannot be successfully built using Zone OH parameters due to serious thermal deformations.Laser Power=120WD R y t i s o r o Pr e w o P r e s a LScanning Speed (mm/s)Laser Power=80WScanning SpeedD R y t i s o r o PScanning Speed (mm/s) Fig. 1 Process Window of SLM and Porosity Distribution at 120W and 80WIn previous research, it was found that specimens can be built with various porosity distribution when using processing parameters, such as laser power and scanning speed, from different melting zones. Thus, investigations about defect morphology are primarily based on specimens when the laser power is equal to 80W and 120W, as shown in Fig. 1.

Thus, investigations about defect morphology are primarily based on specimens when the laser power is equal to 80W and 120W, as shown in Fig. 1. Zone III processing parameters at elevated scanning speeds are of particular interest in this study because they are able to fabricate specimens with controlled porosity. An EOS M270 Direct Metal Laser Sintering (DMLS) system was utilized for building SLM specimens in this study.An Arcam S400 EBM system was utilized in this study. Changes to beam current and scanning speed independently do not cause porosity in specimens, because the Arcam system dynamically coordinates these two parameters to achieve a steady melt pool size. However, “line offset” and “focus offset” can be varied to make porous specimens. Line offset refers to the distance between two hatch lines.

However, “line offset” and “focus offset” can be varied to make porous specimens. Line offset refers to the distance between two hatch lines. And focus offset is the additional current running through the respective electromagnetic coil and can be translated into an offset of the focal plane from its zero position and thus a change in beam area [6]. Increased line offset results in lower energy densities, which forms voids in Ti-6Al-4V specimens. Increased focus offset increases beam diameter, thus lowering energy density and generating porosity.Fig. 2 shows porosity distribution by varying line offset. It can be seen that the porosity value increases when the line offset value is increased, especially when the line offset is beyond 0.18mm. As shown in Fig. 3, apparent porosity (> 0.5%) appears when focus offset is larger than 16mA. Particularly at 20mA and 24mA, the increased focus offset causes a sharp increase in porosity in Ti-6Al-4V specimens.

3, apparent porosity (> 0.5%) appears when focus offset is larger than 16mA. Particularly at 20mA and 24mA, the increased focus offset causes a sharp increase in porosity in Ti-6Al-4V specimens. In order to investigate the defect characteristics, EBM specimens were built by increasing line offset and focus offset. In SLM and EBM processes, a number of processing parameters can be utilized to create specimens with a certain amounts of porosity, such as Zone II and III parameters of SLM or increased line offset and focus offset of EBM. These parameters are called “marginal parameters” in this study.Fig.2 EBM Porosity versus Line Offset Fig.3 EBM Porosity versus Focus Offset2.3 Image Processing MethodFor each marginal parameter combination, two cubical Ti-6Al-4V specimens were sectioned and polished for microscopy. One specimen was cut along the horizontal plane; the other along vertical plane, as shown in Fig. 4.

One specimen was cut along the horizontal plane; the other along vertical plane, as shown in Fig. 4. Cross sections were polished and then observed under optical microscope. Micrographs were taken for image processing process using Matlab R2012a.SLMEBMSpecimen10m mVertical PlaneHorizontal PlaneFig. 4 Ti-6Al-4V SpecimenMicrographs were converted to binary images in order to highlight the features of the defective region. When a pixel’s luminance of the original image is less than a user-selected threshold value, the pixel will be replaced with a value of 1 and other pixels will be replaced with a value of 0. In this paper, the threshold value is selected as 0.2. By doing this, the solid region is shown as a black color, while defective areas are white. Images which included contamination or polishing scratches were “cleaned” using a de-noising algorithm. For each defect, an area value could be obtained by counting pixels and then multiplying by individual pixel area.

For each defect, an area value could be obtained by counting pixels and then multiplying by individual pixel area. All defect areas can be summed up and then divided by cross sectional area to calculate image processing porosity (RI). Certain descriptors such as circularity, convexity, and elongation are also employed for characterizing defect morphology in this study according to:Circularity=24AP(1)Convexity=convexhullPP(2)Elongation=minormajor1LL(3)where A is area, P is actual perimeter, Pconvexhull is the convex hull perimeter, Lminor is the minor axis length, and Lmajor is the major axis length. A perfect circle has a circularity of 1 while an irregular object has a circularity value closer to 0. Convexity is a measure of the surface roughness of a defect. A smooth shape has a convexity of 1 as the convex hull perimeter is exactly the same as the actual perimeter.

Convexity is a measure of the surface roughness of a defect. A smooth shape has a convexity of 1 as the convex hull perimeter is exactly the same as the actual perimeter. A very irregular object has a convexity closer to 0 as the actual perimeter is greater than the convex hull perimeter due to the fine surface features. The major axis length of the defective area refers to the length of the major axis of the ellipse that has the same normalized second central moments as the region. Similarly, the minor axis length specifies the length of the minor axis. A shape symmetrical in all axes such as a circle or square will have an elongation value of 0 whereas shapes with large aspect ratios will have an elongation closer to 1.3. Results and DiscussionStatistical distributions of the morphological descriptors of defects in each cross section are illustrated by histograms. Characteristics of defect morphology are analyzed and compared according to these distributions.

Characteristics of defect morphology are analyzed and compared according to these distributions. Melt pool variation is discussed to explain defect formation.3.1 Characterization of SLM Defects3.1.1 Defect Morphology of Zone II DefectsWhen laser power, hatch spacing and layer thickness are constant, increasing the scanning speed will lower the energy density during SLM [2]. Fig. 5 shows horizontal and vertical cross sections of Ti-6Al-4V specimens built using Zone II marginal parameters, with scanning speed ranging from 360mm/s to 600mm/s.120W 360mm/s Horizontal120W 480mm/s Horizontal120W 600mm/s Horizontal0.5mm0.5mm0.5mm120W 360mm/s Vertical120W 360mm/s Vertical120W 600mm/s Vertical0.5mm0.5mm0.5mmFig. 5 Cross Section of SLM-produced Zone II SpecimensIt can be seen that the defect amount decreas gradually when increasing the scanning speed to the fully dense melting zone. Defects are randomly spread on the horizontal and vertical cross sections.

Defects are randomly spread on the horizontal and vertical cross sections. The defect dimension is also reduced with decreasing porosity. This trend is verified by the histograms in Fig. 6 (a) by the defects’ area and their frequency of occurrence. Large defects only appears with higher porosity. Zone II defects are typically round in shape. Their morphology is well interpreted by the histograms of circularity, convexity, and elongation, as shown in Fig. 6 (b), (c) and (d).RD=5.8% RI=5.9%RD=2.0% RI=1.6%RD=0.5% RI=0.2%λ=6.43 k=0.84λ=7.79 k=0.87λ=6.87 k=0.86RD=5.8% RI=6.2%RD=2.0% RI=2.1%RD=0.5% RI=0.5%λ=6.86 k=0.85λ=4.92 k=0.71λ=8.13 k=0.82(a) Histograms of Defect Area (×103 µm 2) (b) Histograms of Defect Circularityλ=1.69 k=0.31λ=1.83 k=0.31λ=1.49 k=0.31λ=2.09 k=0.35λ=2.82 k=0.44λ=1.97 k=0.37(c) Histograms of Defect Convexity (d) Histograms of Defect Elongation Fig.

6 Distribution of Defect Morphology of SLM-produced Zone II SpecimenAs shown in Fig. 6 (b), a fitting curve shows a Weibull distribution with right skewness for defect circularity. Most Zone II defects have circularity values close to 1, which means they are circular or near-circular. Moreover, all Zone II defects have a convexity value larger than 0.5, as shown in Fig. 6 (c). A circular contour is a typical convex shape. Thus, it is easy to understand why most Zone II defects have a convexity value close to 1. As shown in Fig. 6 (d), the fitting curve of Zone II defect elongation also shows a Weibull distribution, but with different shape compared to circularity. Their left skewness indicates that the minor axis length is close to the major axis length, but not completely equal. According to the circularity and elongation distribution, it can be concluded that most defects have an elliptical contour for Zone II specimens.

According to the circularity and elongation distribution, it can be concluded that most defects have an elliptical contour for Zone II specimens. The shape parameter λ and scale parameter k of the Weibull fitting curve show slight difference between horizontal and vertical cross sections. This implies that the elliptical contours of Zone II defects are distinguishable based upon orientation and thus the material has porosity anisotropy. The λ value of Zone II defect elongation is closer to 0 in the horizontal direction than in the vertical direction. This demonstrates that defects are more circular in the plane parallel to scan surface.3.1.2 Defect Morphology of Zone III DefectsCross sections of Ti-6Al-4V specimens built by Zone III marginal parameters are shown in Fig. 7. Laser power is kept constant at 80W. Scanning speed ranges from 840mm/s to 1320mm/s.

7. Laser power is kept constant at 80W. Scanning speed ranges from 840mm/s to 1320mm/s. When the scanning speed is far away from Zone I, the energy density will be lowered which results in incomplete melting during the SLM process. It is noted that the porosity increased with increasing scanning speed.80W 840mm/s Horizontal80W 1080mm/s Horizontal80W 1320mm/s Horizontal0.5mm0.5mm0.5mm80W 840mm/s Vertical80W 1080mm/s Vertical80W 1320mm/s Vertical0.5mm0.5mm0.5mmFig. 7 Cross Section of SLM-produced Zone III SpecimensDefect dimension increases when increasing scanning speed. Similar to Zone II defects, the higher the porosity, the larger the defect dimension. However, unlike Zone II defects, Zone III defects are in irregular shape, in both horizontal and vertical cross sections. Large defects appears in Zone III specimens but with a small amount, as shown in Fig.

Only a few Zone III defects are circular. Most of them are very irregular. The normal distributions of elongation have mean values ranging from 0.39 to 0.47, with standard deviation from 0.16 to 0.19. It can be seen that more than 70% of Zone III defects have an elongation value less than 0.5. Only 4% of Zone III defects have elongation value larger than 0.7, which means a long and narrow contour. As shown in Fig. 8 (c), some Zone III defects show concave characteristics, which illustrates their irregular shape.3.1.3 Comparison of Defect Morphology in Zone II and Zone III SpecimensGenerally, Zone II and Zone III defects show quite different morphological characteristics, especially in circularity and elongation. In order to further characterize the morphology, specimens with Zone II and Zone III defects are compared in Fig. 9. Both specimens have similar porosity (~5.8%).

In order to further characterize the morphology, specimens with Zone II and Zone III defects are compared in Fig. 9. Both specimens have similar porosity (~5.8%). It is notable that the energy density of marginal parameters (120W&360mm/s) in Zone II is higher than that of Zone III marginal parameters (120W&1560mm/s).Most Zone II defects are round in shape, while Zone III defects are large with an irregular shape. From Fig. 10 (a), it can be seen that most Zone II defects are have an area less than 6×103µm 2, while some Zone III defects have areas up to 14×103µm 2.Weibull distribution is a proper choice to describe Zone II defect morphology, while Zone III defect morphology can be described using a normal distribution.

10 Distribution of Defect Morphology of SLM-produced Zone II and III (Porosity≈5.8%)Zone III defects are formed between melt pools or hatch lines due to insufficient energy density. Thus, it makes sense that Zone III defects are spread stochastically with irregular shape. Both circularity and elongation histograms of defects are normally distributed. An SLM Ti-6Al- 4V top surface fabricated with marginal parameters is shown in Fig. 11.Fig. 11 Top Surfaces of SLM Specimens by Marginal Parameters (a) Zone II and (b) Zone IIIWhen Ti-6Al-4V powder was melted using Zone II marginal parameters, the melt pool was extremely enlarged. Large melt pool results in a large overlap between hatch lines. However, no voids or defects were directly formed by the large melt pool and highly overlapped hatch lines. Spherical pits, as shown in Fig. 11 (a), are believed to cause spherical defects in Zone II specimens.

Spherical pits, as shown in Fig. 11 (a), are believed to cause spherical defects in Zone II specimens. These pits result from the recoating blade scraping particles, which are solidified from the ejected molten materials on the surface. This phenomenon is dramatic because Zone II marginal parameters melt the powder with high energy density. By contrast, the Zone III voids or pores directly generated on the top surface, as shown in Fig. 11 (b), are entrapped by new recoated and melted layers.3.2 Characterization of EBM DefectsEBM defects resulted from the increased line offset and focus offset values. Energy density is decreased due to these marginal parameters. Compared to laser melting, EBM defects are formed similarly to Zone III defects in SLM.

Energy density is decreased due to these marginal parameters. Compared to laser melting, EBM defects are formed similarly to Zone III defects in SLM. In this section, the morphology of line offset defects and focus offset defects are analyzed.3.2.1 Defect Morphology of Line Offset (LO) DefectsPorosity dramatically increases when the line offset value is larger than 0.18mm, according to the porosity distribution curve in Fig. 2. Thus, Fig. 12 shows the cross sections of specimens at specific line offsets of 0.18mm, 0.22mm, 0.26mm, and 0.30mm.LO=0.18mm HorizontalLO=0.22mm HorizontalLO=0.26mm HorizontalLO=0.30mm Horizontal2 mm2 mm2 mm2 mmLO=0.18mm VerticalLO=0.22mm VerticalLO=0.26mm VerticalLO=0.30mm Vertical2 mm2 mm2 mm2 mmFig.

Large defects can be observed directly with bare eyes. It can be seen that, when line offset is 0.18mm there are only a few defects showing on the cross sections. However, as line offset increases, the defect amount increases correspondingly. As shown in Fig. 13 (a), most LO defect areas are smaller than 30×103 µm 2. Large defects are rare, but can have extremely large area. A Weibull fitting curve can be used to describe the distribution of the histograms of LO defect circularity. But the curve characteristics are very different from that of SLM Zone II or Zone III defects. From Fig. 13 (b), it is noted that all the fitting curves show left skewness, which means no defects have a circular contour. Thus, most EBM defects are very irregular.Defect convexity, shown in Fig. 13 (c), verifies the irregularity of LO defects. Although many defects have a convexity value close to 1, there are a large amount of defects with small convexity values.

13 (c), verifies the irregularity of LO defects. Although many defects have a convexity value close to 1, there are a large amount of defects with small convexity values. This indicates that these defects have concave contour, as shown on the cross sections. Elongation of LO defects can be described by normal distribution fitting curves. All fitting curves, as shown in Fig. 13 (d), have mean values close to 0.50, and standard deviation close to 0.20. This is similar to Zone III defects of SLM specimens. Most LO defects (about 84%) have an elongation value less than 0.7.3.2.2 Defect Morphology of Focus Offset (FO) DefectsCross sections of Ti-6Al-4V specimens with FO defects are shown in Fig. 14. Large defects can also be observed with the naked eye from sectioned surface. Some defects are successive on the vertical cross sections. These defects are spread vertically from the bottom to the top surface.

Some defects are successive on the vertical cross sections. These defects are spread vertically from the bottom to the top surface. It is believed that an increased focus offset value generates an extended melt pool. But the reduced melting depth creates an unstable melt pool due to the underneath un-melted powder. Voids and pores are generated due to the instability of the melt pool. This means that FO defects are often created above the voids or pores of a previous layer. Thus, successive defects grows as shown in the vertical cross sections.FO=16mA Horizontal FO=20mA Horizontal FO=24mA Horizontal2 mm2 mm2 mmFO=16mA VerticalFO=20mA VerticalFO=24mA VerticalSuccessive DefectsSuccessive Defects2 mm2 mm2 mmFig. 14 Cross Sections of EBM-produced Specimen at Various Focus Offset ValuesAs for the morphology of FO defects, it can be seen, in Fig. 15 (a), that most defects are small, with an area less than 15×103 µm 2. Only a few large defects were found.

15 (a), that most defects are small, with an area less than 15×103 µm 2. Only a few large defects were found. The circularitydistribution of FO defects can also be described by Weibull fitting curves, as shown in Fig. 15 (b). Few defects show circular contour according to the left skewness of the distribution characteristics. Fig. 15 (c) shows the convexity of FO defects. Most defects have a concave contour. It is predictable that narrow and long defects exists in the specimens. A normal distribution can be used to describe the elongation of FO defects.

15 Distribution of Focus Offset Defect Morphology of EBM-produced Specimens3.2.3 Comparison between LO Defects and FO DefectsThe morphological characteristics of LO defects and FO defects are similar, based on the defect area, circularity, convexity, and elongation distribution of histograms. Thus, it can be determined that the morphology of LO defects and FO defects are the same. The only difference is the spreading characteristics of defects in the specimens. From the cross section of EBM specimens, it can be seen that LO defects are randomly spread on the horizontal and vertical planes. It is hard to say any of defects are correlated with each other. However, successive FO defects are aligned in the vertical plane. The defects resulting from an increased focus offset more easily lead to the growing of defective region.The formation of EBM defects can be further explained by the melt pool characteristics on the top surface.

The melt pools are isolated from each other due to the increased line offset. Molten powder is easily fused together causing voids, as shown in Fig. 16 (a). Although melt pools are highly overlapped when increasing focus offset, a certain amount of powder underneath is not melted due to the shallow depth of the melt pools. The un-melted powder has an impact on the upper surface where voids are formed, as shown in Fig. 16 (b).(a)(b)PoresPores0.5mm0.5mmFig. 16 Top Surface of EBM Specimens (a) LO=0.26mm and (b) FO=20mA3.3 Image Processing Porosity RIThe porosity RD is obtained by measuring the specimen’s relative density. It is considered an accurate method for estimating the porosity of SLM and EBM materials. RI is used as a reference for estimating a specimen’s porosity according to the ratio of defective area to the overall area of each cross section. It is notable that RI is very close to RD.

RI is used as a reference for estimating a specimen’s porosity according to the ratio of defective area to the overall area of each cross section. It is notable that RI is very close to RD. It can be inferred that, if enough cross sections are processed, the image processing porosity RI should be a feasible way of estimating the actual porosity of specimens. Although RI has to be obtained from destructive characterization techniques, it enables the avoidance of lacquering or oil impregnation, which must be applied to the surface of porous parts to avoid penetration of fluid when using Archimedes method. Thus, image processing porosity can be considered an alternative way of estimating porosity of additively manufactured specimens.4. ConclusionDefect morphology of SLM and EBM specimens were studied using image processing to investigate defect area, porosity and other morphological characteristics such as circularity, convexity, and elongation.

It can be concluded that higher porosity is usually accompanied with larger defect dimensions in both SLM and EBM specimens. Defect morphology can be correlated with melt pool characteristics.Zone II defects formed by high energy density are distributed in SLM specimens and have a spherical shape. Porosity is a result of mechanically scraping the solidified particles, which are ejected from melt pool. A Weibull distribution can be used to describe the circularity and elongation of Zone II defects. Zone III defects, which are formed by low energy density levels, are irregular in shape in SLM. Defects are directly generated due to the insufficient energy for complete melting. Porosity is entrapped when new recoated and melted layers are applied over top. Circularity and elongation of Zone III defects are normally distributed.Line offset defects and focus offset defects in EBM are all formed due to a reduced energy density, similarly to Zone III defects in SLM.

Their irregularity can be attributed to the large pores formed on the top surface. There are nearly no circular defects in EBM specimens according to the circularity distribution. A certain amount of defects show concave characteristics.Image processing porosity RI can reflect the approximate porosity of porous parts. Thus, image processing can be a feasible way of estimating porosity if cross section images are well processed.AcknowledgementsThe authors gratefully acknowledge the support of the Office of Naval Research, awards N00014‐09‐1‐0147 and N00014‐10‐1‐0800, Technical Monitor: Dr. Ignacio Perez. The authors also express their gratitude to the staff of the Rapid Prototyping Center at the University of Louisville for their assistance.Reference [1] I. Gibson, D.W. Rosen, B. Stucker. AdditiveManufacturing Technologies: Rapid Prototyping to Direct Digital Manufacturing. Springer, New York, NY, 2009[2] L. Thijs, F. Verhaeghe, T. Craeghs, et al.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDefects Classification of Laser Metal Deposition Using Acoustic Emission SensorHaythem Gaja*, Frank Liou*+ * Mechanical Engineering Missouri University of Science and Technology, Rolla, MO 65409AbstractLaser metal deposition (LMD) is an advanced additive manufacturing (AM) process used to build or repair metal parts layer by layer for a range of different applications. Any presence of deposition defects in the part produced causes change in the mechanical properties and might cause failure to the part. In this work, defects monitoring system was proposed to detect and classify defects in real time using an acoustic emission (AE) sensor and an unsupervised pattern recognition analysis.

In this work, defects monitoring system was proposed to detect and classify defects in real time using an acoustic emission (AE) sensor and an unsupervised pattern recognition analysis. Time domain and frequency domain, and relevant descriptors were used in the classification process to improve the characterization and the discrimination of the defects sources. The methodology was found to be efficient in distinguishing two types of signals that represent two kinds of defects. A cluster analysis of AE data is achieved and the resulting clusters correlated with the defects sources during laser metal deposition.Keywords: Laser metal deposition, Acoustic emission, Deposition defects, Clustering analysisINTRODUCTIONIn general additive manufacturing is extensively used even though monitoring anddetection of defects during AM still require a better understanding.

One of the difficulties in using an adaptive control and LMD monitoring system is the accurate detection of defects as being formed during the metal deposition. The objective of monitoring laser metal deposition process is to prevent and detect damage of produced part for any deposition path and part design. In the LMD process, particular changes in the acoustic emission signal indicate the present of defects, these changes must be carefully considered to ensure the effectiveness of the control system. AE has the advantage of real-time, continuous monitoring of LMD. The central goal of such a system is to indicate the occurrence of defects events, but classifying the type of defect is also necessary for the better use of the system and suggestion of corrective remedies.Bohemen [1] demonstrated that martensite formation during gas tungsten arc (GTA) welding of steel 42CrMo4 can be monitored by means of AE.

It was shown that a particular relation exists between the root mean square (RMS) value of the measured AE and the volume rate of the martensite formation during GTA welding. Grad et al. [2] examined the acoustic waves generated during short circuit gas metal arc welding process. It was found that the acoustic method could be used to assess welding process stability and to detect the severe discrepancies in arc behavior.Yang [3] used an Acoustic emission (AE) sensor to identify damage detection in metallicmaterials. Results suggested a strong correlation between AE features, i.e., RMS value of the1952reconstructed acoustic emission signal, and surface burn, residual stress value, as well as hardness of steels.

Diego-Vallejo [4] in his work found that the focus position, as an important parameter in the laser material interactions, changes the dynamics and geometric profile of the machined surface and the statistical properties of measured AE signal.Recently, Siracusano [5] propose a framework based on the Hilbert–Huang Transform forthe evaluation of material damages, this framework facilitates the systematic employment of both established and promising analysis criteria, and provides unsupervised tools to achieve an accurate classification of the fracture type. Bianchi [6] suggested a wavelet packet decomposition within the framework of multiresolution analysis theory is considered to analyze acoustic emission signals to investigate the failure of rail-wheel contact under fatigue and wear study.

The application was shown to be an adequate for analyzing such signals and filtering out their noise real time monitoring.However, more research needs to be done regarding using the acoustic emission sensor inmonitoring laser metal deposition. In this paper, the defects type distinguishing of the LMD and its corresponding key features are investigated by clustering the AE signals. The acoustic emission (AE) technique is suitable to examine the defeats sources during LMD because of containing rich defect-related information such as crack and pore formation, nucleation and propagation.

Information on defects development is difficult to obtain by only using the AE waveform in a time-space, as a non-stationary process, thus other features such as amplitude, energy, rise time, count and frequency are extracted to analyze qualitatively defects mechanisms.The purpose of the present work is, first to detect laser metal deposition defects as formed layer-by-layer to take the necessary correction action such as machining and remitting, second to develop a reliable method of analysis of AE data during LMD when several AE sources activated to categorize the defects into clusters based on the defect type.EXPERIMENTAL SETUPFigure 1 shows a schematic diagram of the experimental set-up. The YAG laser wasattached to a 5-Axis vertical computer numerical control machine that is used for post-process machining after LMD.

The YAG laser wasattached to a 5-Axis vertical computer numerical control machine that is used for post-process machining after LMD. Picoscope 2205A works as a dual-channel oscilloscope to capture the AE signal and stream it to a computer for further analysis, the oscilloscope measures the change in the acoustic emission signal over time, and helps in displaying the signal as a waveform in a graph. An acoustic emission sensor (Kistler 8152B211) captured a high-frequency signal. The bandwidth of the AE sensor was 100 kHz to 1000 kHz. The raw signals were first fed through the data acquisition system and then processed and recorded using Matlab software.1953Figure 1. Experimental Setup Shown the LMD System and AE Data Acquisition SystemThe acoustic emission signal was recorded during a laser deposition process in an oxidized environment and contaminated powder to induce pores and cracks as a result of thermal coefficient difference. The material of the substrate was tool steel.

The material of the substrate was tool steel. Cracks and porosities were simulated by mixing the mainly Ti-6Al-4V powder with H13 tool steel powder. The two powders particles as illustrated in Figure 2 are non-uniform in shape and size and may contain internal voids which can cause deposition defects when they mixed.(a) Ti-6Al-4V Metal Powder(b) H13 Metal PowderFigure 2 a-b Optical Image of the Metal Powders Used in Deposition ProcessFigure 3 illustrate the main steps in the developed procedure which used to analyze theAE data. A layer is created by injecting the metal powder into a laser beam which is used to melt the surface of a substrate and create a small molten pool and generate a deposit. The AE sensor is attached to a substrate to transform the energy released by the laser deposition into acoustic emission signal. The total length of the deposition is 15 mm was performed with standard parameters for depositing titanium powder as shown in Table 1.1954Figure 3.

The total length of the deposition is 15 mm was performed with standard parameters for depositing titanium powder as shown in Table 1.1954Figure 3. Step-by-step Operations Used to Perform the Acoustic Emission AnalysesTable 1 Laser Metal Deposition process parametersParameter Laser power Powder feed rate Table velocity Length of track Layer thickness Layer widthValue 1,000 Watt 10 g/min 300 mm/min 15 mm About 0.5 mm About 2.5 mmThe AE events can be represented in the frequency domain using Fast Fourier Transform(FFT) or in the time domain using peak amplitude, kurtosis, energy, the number of counts, duration, and rise time. Among all the features, the signal amplitude alone was measured in real time by the data acquisition system. All the other descriptors were calculated from the waveforms at the end of deposited layer because they are very dependent on the amplitude threshold used to detect the arrival time and the end of an AE signal.

Figure 4 shows some of the1955time-dependent features. In this work, all these features were used in a multi-parameter statistical analysis and clustering analysisFigure 4 Time dependent AE event featuresCLUSTERING ANALYSISIt is not easy to discriminate precisely the AE signal associated with each defect source from the waveform of the signal; thus, it is useful to use clustering analysis. Clustering analysis is a machine learning technique which group the AE events based on their features to create clusters in such a way that the AE events inside a cluster are similar to each other, and also dissimilar from events in other clusters. In this work, the K-means clustering algorithm was used to group the AE events into homogeneous subgroups (clusters).

In this work, the K-means clustering algorithm was used to group the AE events into homogeneous subgroups (clusters). A silhouette width value was used to find the optimal number of clusters.The K-means clustering algorithm aims to minimize within-cluster distances between allthe vectors of a cluster and its center and maximize the distances between the centers of all clusters. The clustering algorithm requires the number of clusters k to be known and specified in advance; thus, the silhouette width was used for a range of clusters from two clusters to ten clusters. The number of clusters with the maximum average silhouette width was used to group the AE events into sub-groups reflecting the number of defects. The k-means algorithm can then be described as follows:1. Specify the maximum number of clusters (r). 2. Assume the number of clusters k from 1 to r and randomly initialize each cluster centerCi, where i is from 1 to k3.

Specify the maximum number of clusters (r). 2. Assume the number of clusters k from 1 to r and randomly initialize each cluster centerCi, where i is from 1 to k3. Calculate the Euclidean distance between the vector and the centers of the clusters andthen assign each input vector (or pattern) to the nearest cluster.3. Recalculate the location of the cluster center according to the nearest mean. 4. Repeat steps 2 and 3 until there are no changes in these center locations.19565. calculate the maximum average silhouette width. 6. Repeat steps from 2 to 5 for all possible number of clusters. The greater the silhouette value, the better the clustering results [7, 8].

calculate the maximum average silhouette width. 6. Repeat steps from 2 to 5 for all possible number of clusters. The greater the silhouette value, the better the clustering results [7, 8]. The optimal valueof k is determined according to the maximum of the silhouette width defined as𝑠(𝑘) =1𝑛∑𝑛 𝑙=1𝑚𝑖𝑛(𝑏(𝑙,𝑘)−𝑎(𝑙))𝑚𝑎𝑥(𝑎(𝑙),𝑚𝑖𝑛(𝑏(𝑙,𝑘)))(7)Where 𝑎(𝑙) is the average distance between l-th event and all other events in the same cluster, and 𝑏(𝑙, 𝑘) is the minimum of the average distances between the l-th event and all the event in each other cluster. The silhouette width values range from -1 to 1. If the silhouette width value for an event is about zero, it means that that the event could be assigned to another cluster. If the silhouette width value is close to -1, it means that the event is need to be assigned to other cluster. If the silhouette width values are close to 1, it means that the event is well clustered.

If the silhouette width values are close to 1, it means that the event is well clustered. A clustering can be evaluated by the average silhouette width 𝑠(𝑘) of individual events.The largest average silhouette width, over different K, indicates the best number of clusters. As can be seen, in Table 2 the greatest average silhouette width is 0.8108. The number of clusters was confirmed using Bayesian Information Criterion (BIC) [9], since it is recommended to validate number of clusters through use of several methods. The number of clusters with the lowest BIC is preferred which means k=2 is the optimum number of clusters.Table 2.

The number of clusters with the lowest BIC is preferred which means k=2 is the optimum number of clusters.Table 2. Average Silhouette Width for Different Number of ClustersNumber of clusters Average Silhouette Width2 3 4 5 6 7 8 9 10 110.8108 0.4730 0.4963 0.4770 0.5405 0.5004 0.4387 0.5210 0.4254 0.5346BIC 4409.199532 5307.464994 6268.570481 7268.886407 8331.531180 9397.995038 10464.458896 11532.648846 12645.626117 13763.722664The defects in cluster two tend to have more energy, longer duration, slower rise time,large number of counts, higher amplitude, close to the normal distribution with flatter and light tail distribution, and less frequency compared to the defects in cluster one.Analysis of variance (ANOVA) of the cluster centers (Table 3). As can be seen, most ofthe means of clustering features differ significantly across the two clusters, because the null hypothesis (means are equal) is rejected in a case at significant level ≤ 0.05.

The frequency is not significant which means it has a little contribution to the cluster solution. The features with large F value provide the greatest separation between clusters. As the F value increases as the importance of feature increases, this was also illustrated in Figure 5.1957Table 3.

Analysis of variance (ANOVA) of the cluster centers and features importanceRise Time Peak amplitude Duration Kurtosis Number of counts Counts to peak Energy Average frequency Maximum frequency Standard deviation of frequenciesClusterMean Square 30.629 29.703 30.801 3.449 30.043 28.444 31.137 1.184 10.87012.417df 1 1 1 1 1 1 1 1 11ErrorMean Square .153 .180 .149 .930 .170 .216 .139 .995 .718.674df 35 35 35 35 35 35 35 35 3535F199.608 165.109 207.352 3.709 176.523 131.750 224.074 1.190 15.14018.429Sig..000 .000 .000 .062 .000 .000 .000 .283 .000.000(a) Events in cluster one have less energy1958importance3 5 2 9 4 6 1 10 87(b) Events in cluster one have shorter duration(c) Events in cluster one have slower rise time1959(d) Events in cluster one have less number of counts(e) Events in cluster one have lower amplitude Figure 5 a-e comparing the features in clusters one and two the AE EventsDEFECTS TYPES AND OPTICAL MICROSCOPY STUDYFigure 6 shows waveform samples from cluster one and cluster two.

The waveform fromcluster one is quite different from the waveform from cluster two. The dissimilarities in the events features and waveforms found between the two types of signals lead to the conclusion that the source mechanisms are not the same in both cases.1960(a) Waveform Signal Sample from Cluster One(b) Waveform Signal Sample from Cluster TwoFigure 6 a-b Comparison between the Waveforms from both ClustersAfter preparing the surface of deposited metal, the cracks and pores were observed using an optical microscope, the number of cracks to the number of pores strongly correlated with the number of events in cluster two to the number of events in cluster one. Also from the literature [10], [11] the waveform and the features of the acoustic emission signal created by cracks is similar to the events in cluster two.Figure 7 displays cracks caused by thermal stress.

The temperature gradient of thedeposited layer is large in the direction of thickness during laser deposition process, and the thermal expansion coefficients of the two deposition materials are different, which results in that1961the thermal stress at the combining surface of deposition, thus the cracks are formed. It also occurs with powder contamination in the powder feeder [12].Figure 7 Optical image of a transverse cross-sectioned laser deposit showing a crack and gas porosityThe second type of observed defects is pores which have a spherical form and appear inrandom locations not associated with the microstructure as shown in Figure 8. The possible sources of these porosities are surface powder contamination [11], gasses trapped within the powder particles due to the difference in the powder sizes, and an oxidation effect since the oxygen level was high due not using the chamber to stimulate defect formation in this research.

In fact, surface oxides may most likely remain in the solid state in the melting pool and, as such, upset the wetting mechanisms melted the powder and induces voids.Figure 8 Optical image of a transverse cross-sectioned laser deposit showing a gas porosity1962ConclusionsThe AE signal was collected during the LMD in an oxidized environment with mixedmetal powders to stimulate all possible types of defects. Several defects mechanism were activated and detected by AE sensor. K-Means clustering method was implemented to analyze the AE signals and identify defects source mechanisms.The clustering results successfully distinguish two main defects types and their signal characteristics. The number of clusters to be created does not have to be specified in advance, they only depend on the number of defects being created. Porosities produce the AE signals with shorter decay time and less amplitude. The cracks trigger the AE signals with short durations and high amplitudes.

Porosities produce the AE signals with shorter decay time and less amplitude. The cracks trigger the AE signals with short durations and high amplitudes. The signal energy is a crucial feature in identifying the AE defect source mechanisms.The validation of the proposed methodology has been carried out using an optical microscope; it showed the correlation between the number of acoustic events and the number of defects determined by post-test optical microscopy. The numbers of signal events are in each cluster are in agreement with the rough estimations of the associated numbers of defects.1963REFERENCES[1] Van Bohemen, S M C; Hermans, M J M; Den Ouden, G.,” Monitoring of martensiteformation during welding by means of acoustic emission “, Journal of Physics D, Applied Physics (UK) 34.22 (21 Nov. 2001): 3312 3317.[2]L.Grad, J.Grum, I.

Diego-Vallejo, D. Ashkenasi, H.J. Eichler, “Monitoring of Focus Position DuringLaser Processing based on Plasma Emission, Physics Procedia”, Volume 41, 2013[5] Giulio Siracusano, Giovanni Finocchio, “A framework for the damage evaluation ofacoustic emission signals through Hilbert–Huang transform”, Mechanical Systems and Signal Processing, Volume 75, 15 June 2016[6] Davide Bianchi, András Vernes, ‘Wavelet packet transform for detection of single eventsin acoustic emission signals’, Mechanical Systems and Signal Processing, Volumes 64– 65, December 2015, Pages 441-451[7]R. Gutkin, C.J. Green, S. Vangrattanachai, S.T. Pinho, P. Robinson, P.T. Curtis, On acoustic emission for failure investigation in CFRP: Pattern recognition and peak frequency analyses, Mechanical Systems and Signal Processing, Volume 25, Issue 4, May 2011, Pages 1393-1407[8]Peter J.

The different types of defects deteriorate the mechanical performance of L- PBF components, such as fatigue life, to a different extent. However, there is a lack of recognized approaches to classify the defects automatically and accurately in L-PBF components. This work presents a novel hierarchical graph convolutional network (H-GCN) to classify different types of defects by a cascading GCN structure with a low-level feature (e.g., defect features) layer and a high-level feature (e.g., process parameters) layer. Such an H-GCN not only leverages the multi- level information from process parameters and defect features to classify the defects but also explores the impact of process parameters on defect types and features. The H-GCN is evaluated through a case study with X-ray computed tomography (CT) L-PBF defect datasets and compared with several machine learning methods.

The H-GCN is evaluated through a case study with X-ray computed tomography (CT) L-PBF defect datasets and compared with several machine learning methods. H-GCN exhibits an outstanding classification performance with an F1-score of 1.000 and reveals the potential effect of process parameters on three types of defects.Keywords: Laser beam powder bed fusion, X-ray CT, Hierarchical graph convolutional network, Defects classification.IntroductionAs a promising disruptive technology, additive manufacturing (AM), which creates lightweight 3D components with intricate geometries using layer-wise strategy directly from solid digital models, is evolving from rapid prototyping to revolutionary digital manufacturing technology [1-3]. Laser beam powder bed fusion (L-PBF) is one of the metal AM processes that has enormous potential to fabricate parts with complex geometry for functional and structural applications in different industries, such as aerospace and automotive [4, 5].

However, the complex dynamics such as high dynamic molten pool, ultra-high solidification/cooling rate, or large thermal gradient caused by the variability of process parameters like laser power, scanning speed, hatch spacing, and layer thickness [6] will inevitably generate defects in the inner structure of L-PBF components. These different types of defects, such as keyhole porosity, lack of fusion (LoF) pores, and gas entrapped pores (GEP), can serve as crack initiation sites under cyclic loading [7-9] of fatigue life, which can lead to failure of L-PBF components [10-12]. Therefore, it is important to identify such critical defects and prevent crack initiation to reduce the chance of fracture in L-PBF components.1568Moreover, research on defect classification of L-PBF components with new defect characterization has gained tremendous traction.

Defect features, such as volume, surface area, aspect ratio, and sphericity, extracted from X-ray computed tomography (CT) 3D data are of great help to understand different types of defects in L-PBF components. A novel classification method is urgently needed to automatically classify defects of L-PBF components for high accuracy and efficiency with a massive amount of data available [13].

In summary, there are three critical issues to be resolved in this paper: (1) How to utilize multi-level information to explore process parameters' impact on different types of defects in the L-PBF components? (2) How to characterize the property or features of different types of defects in the L-PBF components? (3) How to automatically classify defects of L-PBF components to avoid a complicated manual classification process? The rest of this paper is organized as follows: the proposed H-GCN for defect classification is detailed in Section methodology; further experiment of H-GCN with a real case study on the defect classification of L-PBF components in Section case study; finally, conclusions and future research topics are summarized in the final Section.MethodologyThis work aims to develop a novel hierarchical graph convolutional network (H-GCN) for automatic defects classification in L-PBF components and quantify the impact of process parameters on the generation of different types of defects in L-PBF components.

This model incorporates process parameters into a GCN unit on its high level and then combines the defect features extracted by content-based image retrieval (CBIR) technique from 3D X-ray CT data and output of the high-level GCN unit into a low-level GCN to classify three types of defects, i.e., keyhole porosity, LoF pores, and GEP of L-PBF components. Meanwhile, the hierarchical structure can also shed some light on the impact of process parameters on the generation of different defect types. The framework of the overall research methodology is summarized in Figure 1.Figure 1. Overall method of H-GCN for defect classification of L-PBF components.1569Defect features extractionContent-based image retrieval (CBIR) is one of the computer vision techniques which can characterize defects on 2D fractography images [14, 15] and 3D X-ray CT scans. All the extracted defect features are summarized in Table 1 with their corresponding descriptions.

All the extracted defect features are summarized in Table 1 with their corresponding descriptions. Apart from traditional features such as volume, surface area, and major axis length, other "untraditional" features can also reveal defect characteristics. For example, solidity can be used to measure whether the shape of a defect is convex and compact with an irregular shape; sphericity is the ratio of the surface area of a sphere to the surface area of the particle [16]; sparseness is the ratio between the volume of ellipsoid and volume of the object. The illustration of defect features for keyhole, LoF, and GEP can be seen in Figure 2. These defect features can be used as one of the input components of the proposed H-GCN in this study.Table 1.

The illustration of defect features for keyhole, LoF, and GEP can be seen in Figure 2. These defect features can be used as one of the input components of the proposed H-GCN in this study.Table 1. Defect features from X-ray CT data of L-PBF components.RepresentationsDefinitionBasics for defect featuresMinor axis length Major axis length Length along z-axis Convex areaVolume Surface area Max axis lengthAspect ratio𝑎𝑎 𝑏𝑏 𝑐𝑐 𝑆𝑆𝑐𝑐 𝑉𝑉𝑉𝑉 𝑆𝑆𝑎𝑎 𝐿𝐿 𝑏𝑏 𝑎𝑎SphericitySolidity𝑇𝑇ℎ𝑉𝑉 𝑠𝑠𝑠𝑠𝑠𝑠𝑠𝑠𝑎𝑎𝑐𝑐𝑉𝑉 𝑎𝑎𝑠𝑠𝑉𝑉𝑎𝑎 𝑜𝑜𝑠𝑠 𝑎𝑎 𝑠𝑠𝑠𝑠ℎ𝑉𝑉𝑠𝑠𝑉𝑉 𝑆𝑆 𝑆𝑆𝑎𝑎 𝑆𝑆𝑐𝑐Descriptors (Features)Equivalent diameter33𝑉𝑉𝑉𝑉[17]𝑑𝑑 = 2�4𝜋𝜋RoundnessExtent𝑅𝑅 =4𝑆𝑆𝑎𝑎 2 𝜋𝜋𝑏𝑏Sparseness𝑀𝑀𝑀𝑀𝑀𝑀𝐷𝐷𝑓𝑓ElongationFlatness𝑇𝑇ℎ𝑉𝑉 𝑠𝑠𝑠𝑠𝑠𝑠𝑠𝑠𝑎𝑎𝑐𝑐𝑉𝑉 𝑎𝑎𝑠𝑠𝑉𝑉𝑎𝑎 𝑜𝑜𝑠𝑠 𝑎𝑎 𝑠𝑠𝑠𝑠ℎ𝑉𝑉𝑠𝑠𝑉𝑉 𝑆𝑆𝑎𝑎 𝑎𝑎 𝑏𝑏𝑏𝑏 𝑐𝑐1570Figure 2.

The procedure and illustrations of parts of defect features extraction from X-ray CT data via CBIR techniques.The basics of graph convolutional networkGraph convolutional network (GCN) is a deep learning model [18], which can combine the feature information and the graph structure to learn better representations of graphs via feature propagation and aggregation [19]. GCN consists of three parts: input layer, hidden layer, and output layer. In the input layer, the initial node representation is to be given. The GCN update the parameters of hidden layers to learn a better node representation and embed over training. The output layer can be used for (1) node classification, where the goal is to learn a representation vector so that the label can be predicted; (2) graph classification, where the set of graphs of an entire graph [20].

Currently, GCN adopts a neighborhood aggregation strategy to iteratively update node representations by {𝐺𝐺1, … , 𝐺𝐺𝑁𝑁} ⊆ 𝒢𝒢 aggregating their neighborhood representations. GCN lays the foundation for investigating classification with the similarity information from features represented by the graph.is used to predict their labels{𝑦𝑦1, … , 𝑦𝑦𝑁𝑁} ⊆ 𝒴𝒴Hierarchical graph convolutional networkHierarchical structure can incorporate various levels of information in the same structure for classification. We propose a hierarchical graph convolutional network (H-GCN) model for classification by a cascading GCN structure with a low-level and high-level features layer. GCN plays a significant role in building up our H-GCN model. Such an H-GCN leverages information from high-level and low-level features for classification and explores the impact of high-level features on low-level features and targets. In this model, we aim to learn representation vectors of a node.

In this model, we aim to learn representation vectors of a node . Graph representation is a technique ℎ𝑣𝑣 to organize data in a graph structure with the relational knowledge of interacting entities [21]. H- GCN adopts the graph data structure to represent the relationship between different entities.such that its label can be classified as𝑦𝑦�𝑣𝑣 = 𝐹𝐹(ℎ𝑣𝑣)𝑣𝑣Definition 1 (Graph): the graph can be denoted as nodes denoted as (the connection 𝑛𝑛 between , it contains a corresponding feature vector is d- as dimensional low-level feature vector. For convenience, we define high-level node feature matrixis the set of, wherewithis the set of edges denoted as, and𝑉𝑉𝐺𝐺 = (𝑉𝑉,  𝐸𝐸).

And the graph𝑣𝑣𝑀𝑀 ∈ 𝑉𝑉 𝑣𝑣𝑖𝑖𝐺𝐺is symmetric. bidirectional, which indicates that the adjacency matrix 𝑉𝑉𝑀𝑀𝑖𝑖 ∈ 𝐸𝐸 𝑣𝑣𝑀𝑀 𝐺𝐺 Assumption 2: Self-loop is allowed for the nodes in graph graph in graph theory [22].which indicates thatis a simple𝐴𝐴𝐺𝐺Definition 2 (Similarity matrix): the similarity matrix similarity between nodes calculated by the Euclidean distance asand. In H-GCN, the similarity matrix𝑀𝑀×𝑀𝑀𝑺𝑺 ∈ ℝ𝐺𝐺is adopted to measure the is𝑣𝑣𝑀𝑀𝑣𝑣𝑖𝑖𝑺𝑺 = �𝑠𝑠𝑀𝑀𝑖𝑖�∀𝑖𝑖, 𝑗𝑗 = 1, … , 𝑛𝑛},(1)2𝑠𝑠𝑀𝑀𝑖𝑖 = �𝑥𝑥𝑀𝑀 − 𝑥𝑥𝑖𝑖�, respectively. For the similarity matrix, where the smaller the distance between two nodes, the stronger the similarity among them. The Min-Max 𝑣𝑣𝑀𝑀 normalization is applied on the similarity matrix to obtain the normalized similarity matrixanddenotes the features of nodesand∀𝑖𝑖, 𝑗𝑗 ∈ {1, 2, … , 𝑛𝑛}𝒙𝒙𝒊𝒊𝒙𝒙𝒋𝒋𝑣𝑣𝑖𝑖to facilitate the determination of edge connectivity.

In among two nodes is given as the mid-point 0.5. Additionally, we where each value on the diagonal is the . In H-GCN model, the adjacency matrix with self- , where and degree matrix is and with low-𝑇𝑇𝑫𝑫 = diag(𝑑𝑑1, … , 𝑑𝑑𝑀𝑀) 𝑑𝑑𝑀𝑀 = ∑ 𝑎𝑎𝑀𝑀𝑖𝑖 = 𝑨𝑨 + 𝑰𝑰𝒏𝒏𝑖𝑖. The adjacency matrix of nodes with high-level features denotes∗ 𝑨𝑨𝑫𝑫� = diag(𝑑𝑑̃1, … , 𝑑𝑑̃𝑀𝑀) .∗ level features is 𝑑𝑑̃𝑀𝑀 = ∑ 𝑎𝑎𝑀𝑀𝑖𝑖 𝑖𝑖∗ 𝑨𝑨𝑯𝑯. Meanwhile, their degree matrixes areand∗ 𝑨𝑨𝑳𝑳The progress of graph representation of H-GCN can be seen in Figure 3. The data is split into high-level features and low-level features and then two normalized similarity matrixes are calculated via the normalized Euclidean distance. Furthermore, adjacency matrix can be obtained . Finally, the high-level and low-level from normalized similarity matrixes based on the threshold graphs are built which can be treated as the inputs of H-GCN.𝑫𝑫� 𝑳𝑳𝑫𝑫� 𝑯𝑯𝑇𝑇1572Figure 3.

Finally, the high-level and low-level from normalized similarity matrixes based on the threshold graphs are built which can be treated as the inputs of H-GCN.𝑫𝑫� 𝑳𝑳𝑫𝑫� 𝑯𝑯𝑇𝑇1572Figure 3. The progress of graph representations of H-GCN model.We define a graph neural network GCN1 with one hidden layer and one neuron in this layer to handle the high-level features to compress features and explore the impact of high-level features on low-level features and different classes of target. For the k-th graph layer in GCN1, we define as the high-level input node representations or embedding matrix (𝑘𝑘−1) of all nodes and as the low-level output node representations. The initial node representations = [ℎ1 𝑯𝑯𝑯𝑯 (𝑘𝑘) can be expressed as 𝑯𝑯𝑯𝑯(𝑘𝑘−1)(𝑘𝑘−1)𝑇𝑇, … , ℎ𝑣𝑣](3)(𝟎𝟎) 𝑯𝑯𝑯𝑯𝑯𝑯which serve as the input to the first GCN1 layer. Then the high-level features are input into the = 𝑿𝑿 hidden layer of GCN1.

In Section data collection, we collect different types of defect data and set the experiment for the L- PBF components. In Section L-PBF defect graph representations, we generate defect graphs via H-GCN based on defects data and process parameters in L-PBF. Finally, in Section analysis of H- GCN for defects classification, we adopt H-GCN for defects classification and evaluate the performance of H-GCN.1574Data collectionIn this case study, the effectiveness of the H-GCN model for defects classification of X- ray CT L-PBF specimens is verified with different L-PBF specimens fabricated with plasma atomized Ti-6Al-4V Grade 5 powder (particle size range of 15 to 53 µm) in an EOS M290 machine supplied by AP&C - a GE Additive company. The geometry of the X-ray CT specimens used in this case study and printed specimens are shown in Figure 4.

The geometry of the X-ray CT specimens used in this case study and printed specimens are shown in Figure 4. In the fabrication process, the process parameters (except layer thickness) were modified from the manufactured recommended values to induce different types of volumetric defects. The EOS recommended process parameters for Ti- 6Al-4V Grade 5 material are 280 W laser power, 1300 mm/sec laser speed, 40 µm layer thickness, and 120 µm hatch distance. The H-GCN model is validated by classifying three different types of defects, i.e., keyhole, LoF, and GEP of L-PBF specimens under eight combinations of process parameters shown in Table 2. Moreover, each specimen contains various defect samples for three types of defects in Table 3.Figure 4. The geometry of the X-ray CT specimens (a) and the printed specimens (b).Table 2.

Moreover, each specimen contains various defect samples for three types of defects in Table 3.Figure 4. The geometry of the X-ray CT specimens (a) and the printed specimens (b).Table 2. The different combinations of process parameters of specimens used in L-PBF experiments.Process parameters Laser powerUnitsWScanning speed mm/s µm Hatch distance µm Layer thicknessS23 224 1300 120 40S16 252 1560 120 40H24 280 1300 132 40SpecimensH23 280 1300 144 40S11 280 1560 120 40K05 336 780 120 40K03 336 910 120 40K15 364 780 120 40It is noted that each defect sample used in this study has 12 size and morphology-related features. The total defect data used in this study is 612 with 97 keyholes, 373 LoF, and 142 GEP. Furthermore, the proportion of training and test set for H-GCN and benchmark methods is 7:3. Additionally, all data are normalized to range by min-max strategy before training and testing.[0, 1]Table 3.

Additionally, all data are normalized to range by min-max strategy before training and testing.[0, 1]Table 3. The statistics for each type of defects for different specimens in L-PBF experiment.Data# of keyhole # of LoF # of GEP TotalTotal number of defects 97 373 142 612S23 - 74 2 76S16 - 92 - 92H24 - 36 3 39SpecimensH23 - 69 7 76S11 - 77 2 79K05 68 25 122 215K03 4 - 6 10K15 25 - - 251575L-PBF defect graph representationsTo build the L-PBF defect graph, we consider two aspects: (1) nodes and (2) edge connections. The L-PBF defects graph is combined with two H-GCN graphs as a high-level consisting of high- and a low-level graph graph level nodes only with three process parameters features and edge’s connection. The edge’s connection is based on the similarity between different nodes.

The edge’s connection is based on the similarity between different nodes. As for the high-level edge’s connection, the similarity of high-level defect nodes with process parameters features is calculated by Equation (1) and shown in Figure 5 (a). Based on this similarity, we apply the threshold of 0.5 to build the connection among high-level defect nodes (similarity < 0.5). The adjacency matrix of high-level defect nodes and their graph representation is shown in Figure 5 (b) and (c). In Figure 5, it is noted that the similarity within keyhole, LoF, and GEP is relatively large (the distance is small) compared with their cross area, which indicates that the defects with process parameters features have a potential relationship. It can also be reflected in its adjacency matrix and graph representation. We observe that almost all high-level defect nodes are connected with the same types of defects and only a few are associated with other types of defects.

We observe that almost all high-level defect nodes are connected with the same types of defects and only a few are associated with other types of defects. For example, the LoF defect nodes (green) in the middle of Figure 5 (c) are connected with LoF, GEP, and keyhole.. We first discussed the high-level graph𝐺𝐺𝐺𝐺𝐻𝐻𝐺𝐺𝐿𝐿𝐺𝐺𝐻𝐻Figure 5. The similarity of defect nodes with high-level features and graph representation. (a) The similarity of defects with high-level features; (b) the adjacency matrix of defect nodes with high-level features; (c) the graph representation of nodes with high-level features.consists of low-level nodes only with 12 defect features and edge connections. Figure 6 (a) and (b) show the similarity of low-level defect nodes and their distributions.

Figure 6 (a) and (b) show the similarity of low-level defect nodes and their distributions. Critically, the threshold of 0.5 cutoff several connections within the keyhole, LoF and GEP similarity based on the results of their distribution which indicates that a few defects show no strong relationship with other the same types of defects.The low-level graph𝐺𝐺𝐿𝐿Figure 6 demonstrates the adjacency matrix and low-level graph representation. It is noted that there are several connections among different types of defects in Figure 6 (a) which is reflected in Figure 6 (b). The number of connections between different types of defects is increased compared with the high-level graph. For example, there are more connections among keyhole, LoF and GEP in Figure 6 (b). The complete graph information for the high-level graph and the low- level graph can be seen in Table 4.1576Figure 6. The similarity of defect nodes with low-level features.

The complete graph information for the high-level graph and the low- level graph can be seen in Table 4.1576Figure 6. The similarity of defect nodes with low-level features. (a) The similarity of defects with low-level features; (b) the distribution of similarity matrix of low-level features; (c) the adjacency matrix of defect nodes with high-level features; (d) the graph representation of nodes with high-level features.Table 4. The graph information of H-GCN for defects classification of L-PBF components.Graph typeH-GCN# Nodes612# Edges180584Number of features for each node 3Average node degree295.07Contains isolated nodes FalseH-GCN𝐺𝐺𝐻𝐻61232257412527.08False𝐺𝐺𝐿𝐿Analysis of H-GCN for defects classificationAs for the evaluation of H-GCN, three statistical metrics precision, recall and F1-score are used for evaluating the model classification performance.

The proposed H-GCN is compared with GCN, neural network (NN), and support vector machine (SVM). The data used for comparison of H-GCN with other benchmark methods are as:H-GCN: adopts defect features + process parameters + similarity among defects. • GCN: adopts defect features + similarity among defects. • NN and SVM: adopt only defect features.The mean and standard deviation of defect classification performance of the H-GCN and other benchmark methods are summarized in Table 5. It is noted that H-GCN significantly outperforms GCN, NN, and SVM for precision, recall and F1-score with all 1.000 without variance. Furthermore, we can observe from the confusion matrix in Figure 7 that NN and SVM are typically failed to classify LoF and GEP with some data shown in Figure 7 (c) and (d),1577respectively.

Furthermore, we can observe from the confusion matrix in Figure 7 that NN and SVM are typically failed to classify LoF and GEP with some data shown in Figure 7 (c) and (d),1577respectively. Moreover, GCN is better than NN and SVM for classifying all three types of defects, especially for GEP, without misclassified data because it uses additional edge connection information. Critically, H-GCN can classify keyhole, LoF, and GEP better than other methods without misclassified data for all types of defects since it adopts not only defect features but also process parameters and their similarity.Table 5.

The defects are as a ratio among typically correlated to the process energy and described by energy density ), and layer thickness ( laser power (), scanning speed (), hatch distance () [23-26]: 𝐸𝐸𝑣𝑣 𝛿𝛿𝑃𝑃𝑣𝑣ℎ,𝑃𝑃the layer thickness is a constant in this case study.𝐸𝐸𝑣𝑣 =𝑣𝑣⋅𝛿𝛿⋅ℎ , [𝐽𝐽\𝑚𝑚𝑚𝑚3]From our experiment for the study of weights of H-GCN, we found that the laser power is ) from laser positively correlated with keyhole since the weight path is power to keyhole. According to the Equation (10), we know that a higher laser power intends to inform keyhole defects.

According to the Equation (10), we know that a higher laser power intends to inform keyhole defects. Additionally, LoF is induced by insufficient energy density defects which(3.65) → (−2.8) → (−4.01578(10)from laser is corresponded with the weight path of experimental result power to LoF and less laser power (low energy density) tends to increase GEP in the inner structure of L-PBF specimens with two weight paths(3.65) → (−2.8) → (3.01)andfrom laser power to GEP.(3.65) → (−0.4) → (4.54)(3.65) → (−2.8) →(0.8)Conclusion and future workThis research develops a novel H-GCN model for defect classification of L-PBF components using a graph representation of data in a layer-by-layer GCN structure to build the multilevel multifactorial model. Specifically, H-GCN first incorporates process parameters and defect features observed from X-ray CT scans into high-level and low-level graphs, respectively, based on the similarity of defects.

Specifically, H-GCN first incorporates process parameters and defect features observed from X-ray CT scans into high-level and low-level graphs, respectively, based on the similarity of defects. The similarity information considers the relationship among defects and improves the accuracy of defect classification. Furthermore, H-GCN adopts two GCNs for high-level and low-level graphs to make highly accurate defect classification. This hierarchical structure of H-GCN enables its flexible scalability. Multilevel information can be processed hierarchically by H-GCN to account for differences and interactions among different level data. This work achieves perfect defect classification of L-PBF components with a statistical F1-score of 1.000.

This work achieves perfect defect classification of L-PBF components with a statistical F1-score of 1.000. Compared with other models such as SVM, its performance improves by about 6%.The future work of H-GCN includes: (1) to extend the hierarchical structure of H-GCN to cases with multilevel multifactorial modeling; (2) to apply H-GCN to other fields and find new applications in a broader range of disciplines, such as drug design in chemistry, cancer subtype classification in biology, and social relationship understanding in social media.References[1][2][3][4][5][6][7][8][9]S. Ford and M. Despeisse, "Additive manufacturing and sustainability: an exploratory study of the advantages and challenges," Journal of cleaner Production, vol. 137, pp. 1573-1587, 2016. N. Guo and M. C. Leu, "Additive manufacturing: technology, applications and research needs," Frontiers of mechanical engineering, vol. 8, no. 3, pp. 215-243, 2013. S. K. Everton, M. Hirsch, P.

T. N. Kipf and M. Welling, "Semi-supervised classification with graph convolutional networks," arXiv preprint arXiv:1609.02907, 2016. F. Scarselli, M. Gori, A. C. Tsoi, M. Hagenbuchner, and G. Monfardini, "The graph neural network model," IEEE transactions on neural networks, vol. 20, no. 1, pp. 61-80, 2008. K. Xu, W. Hu, J. Leskovec, and S. Jegelka, "How powerful are graph neural networks?," arXiv preprint arXiv:1810.00826, 2018.[21] W. L. Hamilton, R. Ying, and J. Leskovec, "Representation learning on graphs: Methods and[22] [23][24][25][26][27][28]applications," arXiv preprint arXiv:1709.05584, 2017. D. B. West, Introduction to graph theory. Prentice hall Upper Saddle River, 2001. H. Gong, K. Rafi, H. Gu, T. Starr, and B. Stucker, "Analysis of defect generation in Ti–6Al–4V parts made using powder bed fusion additive manufacturing processes," Additive Manufacturing, vol. 1, pp. 87-98, 2014. T. Vilaro, C. Colin, and J.-D.

1, pp. 87-98, 2014. T. Vilaro, C. Colin, and J.-D. Bartout, "As-fabricated and heat-treated microstructures of the Ti- 6Al-4V alloy processed by selective laser melting," Metallurgical and materials transactions A, vol. 42, no. 10, pp. 3190-3199, 2011. B. Vrancken, L. Thijs, J.-P. Kruth, and J. Van Humbeeck, "Heat treatment of Ti6Al4V produced by Selective Laser Melting: Microstructure and mechanical properties," Journal of Alloys and Compounds, vol. 541, pp. 177-185, 2012. L. Thijs, F. Verhaeghe, T. Craeghs, J. Van Humbeeck, and J.-P. Kruth, "A study of the microstructural evolution during selective laser melting of Ti–6Al–4V," Acta materialia, vol. 58, no. 9, pp. 3303-3312, 2010. J. Xiong, Z. Xiong, K. Chen, H. Jiang, and M. Zheng, "Graph neural networks for automated de novo drug design," Drug Discovery Today, vol. 26, no. 6, pp. 1382-1393, 2021. S. Rhee, S. Seo, and S.

J. Xiong, Z. Xiong, K. Chen, H. Jiang, and M. Zheng, "Graph neural networks for automated de novo drug design," Drug Discovery Today, vol. 26, no. 6, pp. 1382-1393, 2021. S. Rhee, S. Seo, and S. Kim, "Hybrid approach of relation network and localized graph convolutional filtering for breast cancer subtype classification," arXiv preprint arXiv:1711.05859, 2017. Z. Wang, T. Chen, J. Ren, W. Yu, H. Cheng, and L. Lin, "Deep reasoning with knowledge graph for social relationship understanding," arXiv preprint arXiv:1807.00504, 2018.[29]1580.

flection and the Prevention of Ingress within Laminated Tooling for Pressure Die Casting.R.C.Soar & Dr P.M.Dickens.The University ofNottingham, Centre for Rapid Prototyping, Department ofManufacturing & Operations Management, University Park, Nottingham, UK. NG7 2RD.ABSTRACTWrthin the context of rapid tooling, we are currently assessing the fundamental limitations of laminated tooling for pressure die-casting (PDC) applications. The use of individual laminates to form a die-cast tool presents it own problems, namely the prevention of excessive deflection that may lead to the ingress of pressurised molten aluminium between laminates. Ultimate solutions lie with bonding and clamping techniques of which work is already underway.

Ultimate solutions lie with bonding and clamping techniques of which work is already underway. This paper describes an initial study to establish the fundamental laminated die behaviour in extreme die-casting environments.Keywords: Pressure Die-Casting, Deflection.Laminated Tooling, Rapid Prototyping, Freeform Fabrication, Rapid Tooling,1.INTRODUCTIONOf all of the Rapid Tooling processes, Laminated Tooling is probably the simplest to conceptualise. The process uses the layered data from a 3D CAD model of a tool. Each slice is exported via DXF to a laser profiling machine. Each of the DXF files defines one laminate of the tool and all are nested to :fit a pre-defined sheet of steel, aluminium, stainless steel, etc. After cutting they are de-burred and assembled into the finished tool.

After cutting they are de-burred and assembled into the finished tool. The benefits of laminated tooling1 can be summarised as follows:-The production of large-scale tooling, as the size of each laminate is only restricted by the size ofthe laser profiling bed.The inclusion of conformal cooling channels for decreased cycle times. • The replacement ofdamaged or worn laminates. • The exchange of laminates for different profiles within a tool. • Low cost and time of production, as there is little capital layout due to the abundance oflaser sub-contractors.Laminated tooling is becoming attractive to die-casters because of the huge expense of conventional die production. e.g. Dies commonly require modifications after manufacture; and there is the problem ofreducing hot spots within a die. Conventional die manufacturing may only allows for one attempt to get the design correct, the die must also be dedicated to producing many307tens of thousands of parts to justify the cost.

Conventional die manufacturing may only allows for one attempt to get the design correct, the die must also be dedicated to producing many307tens of thousands of parts to justify the cost. The rapid increase in the use of die-casting, particularly of aluminium to reduce fuel consumption and increase Performance in cars, has In addition, product lines can change annually requiring resuhed in larger dies, nlDning faster. new tooling. Laminated tooling has the potential to offer low cost, large scale dies for limited runs. Even if they cannot Perfectly match the Performance of a conventional die they can allow the die-caster to produce prototype tools that can be run on the die-cast machines. This makes possible the following: the study of material flow throughout the die; the formation of hot spots and soldering (mohen material bonds to the die); the effectiveness of cooling channels; ejector pin layouts; vortices; overflows; gating; etc.

By exchanging laminates within the die many iterations can be carried out before the final die design is set.Some of the groups involved in the development of Laminated tooling around the world include: Stratoconception~, CIRTES (France); Nottingham3 , Warwick, Leeds & Liverpool Universities (UK); MlT\ Clemsons & Ohio State6 Universities (USA); D117 (Denmark); Tokyo University8 (Japan); most are backed by major automotive and aerospace sponsors keen to see a viable process.2.LAMINATED TOOLS FOR PRESSURE DIE-CASTING.For the experimental work an EMBI00 hot and cold chamber pressure die-casting machine was used. This is by no means the largest die-casting machine but it is fuirly typical. Specifications in its cold chamber set up are: -Die Loc1dn2 Force Size ofMoving Platen We4ilit per Shot (Al) Volume per shot Dia. OfIni. Plunger Total Force on Plunger Max. Pressure on Metal Min.

Specifications in its cold chamber set up are: -Die Loc1dn2 Force Size ofMoving Platen We4ilit per Shot (Al) Volume per shot Dia. OfIni. Plunger Total Force on Plunger Max. Pressure on Metal Min. Cycle Time75 Tons' 16" by 16" .651bs 6.5 ins3 1.25" 11,775lbs 9,600 lbs/in2 4secs..~76 Tonneq(MP.tric) 406by406mm .29 kg 106 em3 31.8mm 5,341 kg 674MPa 4 seesThe casting material chosen was Al-Si8-CuJ or LM24 (BSI490)/A380 (ASTM). This alloy is globally used as one of the most applicable to pressure die-casting. Cast aluminium is by :far the largest sector in the PDC field2 • However, it readily oxidises and is aggressive to steel dies. It also has the highest melting point at 750°C. If a laminated die can withstand the pressure die-casting of aluminium it will be suitable for zinc, magnesium as well as low pressure applications.The research took four routes to achieve the aim of viable laminated tooling for pressuredie-casting: -1.

Selection of suitable sheet material and thickness. 2. Testing laminate stacks against failure through thermal cycling/stress.3083. Assessing the fundamentals of laminated die behaviour to withstand deflection andingress ofmolten aluminitlllL4. The bonding of laminates for extreme environments.The first two sections have been covered in previous papers in conjunction with United States Committee for Automotive Research (USCAR)10. A recent paper, in conjunction with GEe Marconi Hirst Divisionll , covers the last section. This paper asks the more fundamental question of whether inter-laminar bonding is required at all.

A recent paper, in conjunction with GEe Marconi Hirst Divisionll , covers the last section. This paper asks the more fundamental question of whether inter-laminar bonding is required at all. Also, what design constraints exist for laminated dies and over what aspect ratio will individual laminates fail or distort.3.PROBLEMS OF LAMINATED DIES IN PDC ENVIRONMENTSWhen using laminated tooling for pressure die-cast applications there is the possibility that molten material may force itself between the laminates of an up-stand feature within the die. If this occurs, then the subsequent casting may solidify onto a feature and be impossible to remove without damaging the tooLFor any laminate die design, the amount of deflection that occurs on an up-stand feature will be proportional to its height, area and geometry, as well as its orientation to the flow of the incoming material. Essentially, for any given sheet material there will be a design limit.

Essentially, for any given sheet material there will be a design limit. The objective of this research is to find out, considering the unique conditions of die-casting, where this limit lies.Due to the dynamic loading on the laminates, as molten material enters the die, the degree of deflection and the amount of ingress can only be established through direct observation of a laminated die in operation. What complicates matters further, is the nature of the molten material as it flows around an up-stand and fills the die. Depending on the influences of the various elements molten material will act as a dynamic load. It is not static, due to the brief time it takes In effect, to :fill the cavity and the constantly changing pressure and velocity within the die. deflection will reach its maximum whilst material passes over the up-stand.

deflection will reach its maximum whilst material passes over the up-stand. However, it will cease as soon as the chamber is full producing an equal pressure on both sides ofthe up-stand.Four things could happen at this point and are illustrated in the Figure 1, overleaf1. Molten material could flow over the up-stand and the pressure could equalise on both sides before any significant deflection or ingress ofmaterial occurs.2. If ingress has occurred in the up-stand during material inlet, the chamber could remain hot enough to allow the equalising pressure to force that material out of the laminate up-stands.3. Any material that forces its way between the laminates could chill so quickly that it will solidify before the back pressure can force it out and so cause the part to freeze onto the ejector side ofthe die set.4. if the deflection is too great the laminate could deform.

if the deflection is too great the laminate could deform. plastically and could normalise into this new position309the When molten material chamber the end laminates will briefly deflect. Whether there is a permanent ingress of material depends on the equalising effect when the material flUs the chamber.entersrills the diechamber As material completely and the pressure equalises, the elasticity in the deflected laminate may allow it to spring back. If the material the laminate is chills too fast whilst deflected, the part will freeze onto the up stand.Figure 1. Effects ofMaterial Flow within Die Cavity.A final point to make about the possible outcomes relates to the viscosity/fluidity and wetability ofthe molten material.

Effects ofMaterial Flow within Die Cavity.A final point to make about the possible outcomes relates to the viscosity/fluidity and wetability ofthe molten material. Even with a large deflection, ingress may not be possible due to the viscous nature ofmolten aluminium.4.SELECTING SUITABLE SHEET MATERIAL.There is generally considered to be only one type of steel that is suitable for aluminium pressure die-casting: This is HI3 hot work tool steel (BHI3 in the US).The selection and location ofsuitable sheet material for this experiment has been a central part ofthe work so far. One ofthe first decisions made was to establish a suitable thickness for the sheet material. The trade-off is between the degree of finish and detail required in the assembled laminate tool against the assembly time and availability of the material. The fastest way to build a laminated tool is to use very thick sheet material.

The fastest way to build a laminated tool is to use very thick sheet material. But, as Figure 2 below shows, if the sheets are too large then all the detail is lost and secondary finishing becomes 90% of the job. The other extreme means the laminates would resemble a metal foil, as in Figure 3. The detail would almost be perfect but hard steels do not come this thin and the tool would be difficult to assemble.Figure 2 & 3. Effects of Laminate Thickness on Steeping310The most readily available sheet thickness for high strength and thermally resistant steels, giving adequate detail, is one millimetre. One millimetre is the thinnest that most sheet materials can be purchased for high performance steels such as H13.

One millimetre is the thinnest that most sheet materials can be purchased for high performance steels such as H13. Rolling steel this thin can improve the grain structure and alloy distribution, but it does set up stresses that must be relieved later on (this is done during final hardening and tempering).5.THE LAMINATED TEST DIE.The ultimate test die design consists of a number of isolated laminate up-stands. Each up stand has a different aspect, that will present one laminate in each group to receive the :full force of the incoming, pressurised, molten aluminium. At a certain height there will be enough force to deflect this laminate to cause an ingress of molten material between itself and the laminate it abutts. Figure 4, below, shows a cross section ofthe two extremes ofthe up-stand arrays that will be created in the test die.Cover DieCover DieEjector DieEjector DieFigure 4.

Figure 4, below, shows a cross section ofthe two extremes ofthe up-stand arrays that will be created in the test die.Cover DieCover DieEjector DieEjector DieFigure 4. Effects ofDeflection on Laminate Upstands.The simplified illustration (A) shows molten material entering from the left and being forced upwards at 450 over a ramp formed by the laminates. Each laminate stands 1 mm. higher than the laminate on the left. This design will not incur any deflection as the laminates in the up stand support each other and the end laminate does not protrude into the flow ofmohen metal.On illustration (B) material enters from the left and is directed up the laminate ramp where it will strike the last laminate before passing over and around it. This laminate will deflect but may move to the upright position again, due to its elasticity/rigidity, before the cast freezes.

This laminate will deflect but may move to the upright position again, due to its elasticity/rigidity, before the cast freezes. Trying to measure this deflection as it occurs in the die would be impossible, this measurement can only be taken by examining the casting after removal from the die. The answer to measuring deflection lies in the second laminate positioned next to the tallest up-stand. If the last laminate deflects enough then there will be an ingress of aluminium that will freeze between the laminates. When the cast is removed from the die a 'witness mark' will remain that can be measured as a direct indication ofthe amount ofdeflection that occurred in the last laminate, as shown in Figure 5.I=:XIMark on Cast I~Figure 5. Measurable Witness Marks311Figures 6, 7 and 8 show the plan and cross-sectional view of the two halves of the laminated die. The laminated test die will be clamped into a bolster.

Measurable Witness Marks311Figures 6, 7 and 8 show the plan and cross-sectional view of the two halves of the laminated die. The laminated test die will be clamped into a bolster. The dimensions are shown below in Figure 9: ..A....- - 246 mm (10") - - - . . <4- 152.4 mm (6") -+20mmT A127-ImIl82mm56mm....- - - - - 296 mm (11.5") ---JoPlan and cross section through the two die balves. Dimensions bave been dictated by the limitations on the EMB 100.Figure 9. Complete Die Assembly.The end-plates will be constructed in H13 to spread the clamping load on the individual laminates and prevent them from warping. Previous testing with GEC and USCAR has shown that a pressure of at least 10 MPa is necessary on each laminate to ensure rigidity and the elimination ofany gaps between laminates this pressure will be provided by the eight MI0 botts.The tool will run for around 500 shots as there is a time constraint on the use of the EMB100 die-casting machine.

Over this time tool wear will be monitored as an indication of potential prototype life.6.DISCUSSION.This tool is now currently in testing and the results will be published early in 1998. The resuhs from this work will dictate where the research will move next. If ingress occurs too readily then the next stage will be to investigate suitable bonding techniques to overcome ingress in these If ingress does not occur readily, then un-bonded laminate dies will be explored applications. further. To produce a laminated tool with no inter-laminar bonding for pressure die-casting would be a huge cost saving over having to use some bonding medium. Instead of looking into bonding techniques, the research would move onto look at clamping techniques.The data collected from this experiment will then be used to formulate a model that will allow the designer of a laminated tool to know the limitations of this process when applied to pressure die-casting.

End plates and a central clamping plate allow the interchange of laminates to change the profile.areshowntoFigure 8.Completed ejector die showing wedges the lasminated die in the bolster. Ejector pins have been removed.rousedhwdFigure 11Laminates that make up the Cover Die assembled and clamped using solid end plates and a solid central core facilitate to disassembly.315Figure 10Two of the four rows of showing laminated arrays the laminate up-stand that will receive the full force of incoming molten aluminium to measure deflection.Figure 12The completed lamjnated test die showing ejector pins and the fan gate in the cover die. The ground parting line makes the laminates hard to distinguish.316.

FORMATION POST-PROCESSING OF ADDITIVE MANUFACTURINGCOMPONENTSN. B. Crane, C. P. Lusk, J. Nussbaum, Y. Consuegra ReyesDepartment of Mechanical Engineering, University of South Florida4202 E. Fowler Ave ENB 118, Tampa, FL 33549AbstractParts produced by additive manufacturing (AM) often require post processing to improvesurface finish and mechanical properties. However, little attention has been given to including deformation in the post processing. Deformation post-processing can address some part size, manufacturing cost, and geometry limitations of 3D printing. Additionally, it could be used to create 3D surfaces using planar manufacturing processes (such as printed circuit board manufacturing). The challenge of deformation post-processing is the design of the correct fabrication state to produce the desired final state and the accurate deformation of the parts to the desired final state.

The challenge of deformation post-processing is the design of the correct fabrication state to produce the desired final state and the accurate deformation of the parts to the desired final state. This paper demonstrates the geometric capability, potential applications, and methods for accurately and repeatedly deforming the initial geometry to the desired configuration using features in the parts.IntroductionAdditive manufacturing (AM) relies on computer control of material bonding/depositionrather than part-specific tooling to define the geometry of the manufactured component. Typically, parts are built up layer by layer until they are complete. After the geometry is formed, the parts are post-processed to achieve the final geometry and properties.

Typically, parts are built up layer by layer until they are complete. After the geometry is formed, the parts are post-processed to achieve the final geometry and properties. The post-processing methods vary with the AM process, but can include removal of support material, infiltration of pores, surface finishing, annealing, and cleaning (Gibson et al., 2010).The AM approach provides substantial freedom in the geometries that are created and eliminates the time requirements of manufacturing specialized tooling. Additionally, some of these processes are capable of integrating multiple materials to create multifunctional components (Wicker and MacDonald 2012, Dollar and Howe 2006, Kataria and Rosen 2001, Malone, 2004). However, the processes themselves are typically much slower than traditional manufacturing methods.

However, the processes themselves are typically much slower than traditional manufacturing methods. Additionally, despite substantial progress, the materials available for AM are more limited than for traditional processes.This paper considers the use of deformation in the post processing of AM components to address some of these challenges. In deformation post-processing a part would be printed in one configuration and then during post-processing deformed into its final configuration. If this approach were applied to printing thin shells (such as what could be produced from sheet metal processing), the printing time and use of support material could be dramatically reduced. As an added benefit, more of the key features could be aligned where printing accuracy and surface finish is often highest.

As an added benefit, more of the key features could be aligned where printing accuracy and surface finish is often highest. (Xu et al., 1999; Flores et al., 2003; Ollison and Berisso, 2010)Deformation post-processing could also be integrated with traditional planarmanufacturing processes to create 3D geometries. Thus, high speed, mature planar cutting and printing processes could be used to rapidly create 3D structures. This would be of particular value in fabricating multifunctional structures incorporating electronics from available printed circuit board technologies.The keys to enabling this alternative approach are:1) Avoid material Failure: Many materials of interest are relatively brittle and so the deformation must be achieved with minimal component strain to avoid failure. 2) Simplify the Deformation process: In the spirit of AM, it is important that the use of specialized deformation tooling be minimized.

2) Simplify the Deformation process: In the spirit of AM, it is important that the use of specialized deformation tooling be minimized. Preferably, the deformation response will be built into the component itself. Of course, the deformation process must be accurate as well in order for the final part geometry to be acceptable.3) Simplify the part design process: In order to be applied to a variety of parts, it must be easy to integrate the required printing features into the components to achieve the desired outcomes.In order to demonstrate the technical feasibility of these methods, this paper will utilize the temperature variation in mechanical properties to avoid material failures. The deformation process will be simplified by reference to both material properties and compliant mechanism design. These concepts are demonstrated for creating 90 degree bends.

The deformation process will be simplified by reference to both material properties and compliant mechanism design. These concepts are demonstrated for creating 90 degree bends. Additionally, a benchmark multifunctional part will be shown.Deformation StrategiesIf the desired post-processing deformation cannot be achieved by custom tooling, then alternative approaches are required. In this work, we consider controlled deformation through localized variation in the material properties and localized variation in the stiffness. Combining these two approaches permits accurate deformation of AM components. These approaches are discussed independently and then examples are presented. All 3D Printed components were produced using a Dimension sst-768 out of ABS.Spatial Variation in Material PropertiesWhile some AM processes are capable of printing variable material, the majority canonly deposit a single structural material per part.

Additionally, since prost processing deformation is a one-time process, the ideal joint material would be flexible for deformation and then stiff during the use phase. The easiest route to achieving this property variation in polymeric AM materials is through heating (Nikzad et al., 2011). This is especially effective in polymer materials where heating above the glass transition temperature significantly reduces the material stiffness and increases the strain to failure.The ideal hinge joint would have no torsional stiffness with infinite axial stiffness withjust one axial degree of freedom. Typically, compliant mechanisms are designed for elastic motion and so this problem is addressed by making a beam very thin in the transverse direction. As bending stiffness varies with the cube of the beam thickness while axial stiffness varies linearly with thickness, this is a feasible approach.

As bending stiffness varies with the cube of the beam thickness while axial stiffness varies linearly with thickness, this is a feasible approach. However, there are obvious compromises in the strength of the resulting element in use.This concept can be integrated with the material properties of polymers to achieve asimilar affect and permit temporary torsional flexibility while maintaining adequate axial stiffness. This can be accomplished by non-uniform heating of the part. The hottest regions will experience a dramatic decrease in their modulus while the cooler regions are relatively unaffected. The easiest way to accomplish this would be to apply heat to the exterior of the part such that the Biot number (Bi =hs/k) is large. Under these conditions, there will be a substantial (though temporary) gradient in the temperature through the part thickness.

Under these conditions, there will be a substantial (though temporary) gradient in the temperature through the part thickness. Since the thickness of the bending joints is much less than the component body, the joints heat through much quicker than the body. This softens the joints and allows the joints to easily undergo deformation while maintaining a much higher stiffness in the component body. Due to the low thermal conductivities of polymers, this condition is more easily achieved than in metals and the processing window will be larger.Figure 1 Demonstration of thermal bending without tensile strain for three different parts. Pairs compare a heated and unheated part of each type.

Pairs compare a heated and unheated part of each type. From left to right, ABS plastic produced by FDM with the curve vertical, ABS plastic made by FDM printed in the position shown, and laser-cut acrylic.To demonstrate the feasibility of bending a part without causing significant axial strain, a small test specimen illustrated in Figure 1 was created from both ABS plastic (FDM) and acrylic (laser cut). The ABS part was printed in two different build orientations. The planar print produced two thin strips of polymer with a gap between while the vertical build produced a filled structure.A weight of 18 g was attached to one end of the specimen while the other was held in afurnace set to 200 C. When the curved segment was observed to straighten, the sample was removed from the furnace and quenched under running water.

When the curved segment was observed to straighten, the sample was removed from the furnace and quenched under running water. The change in the length of the straight segments was within the measurement errors for two part types and modest (1.5%) for the third while the final length of the curved segment was very close to the length of the centerline of the original arc despite a large change in the gap between the rigid blocks and accompanying change in curvature of the thin segment. This demonstrates the potential for bending deformation without significant axial strain.Table 1 Measured strains in test components from heating under tensile load (Strainis the change in length before vs. after the test. Given the difficulties of measuring arc lengths, the nominal arc length was used to calculate strain.

after the test. Given the difficulties of measuring arc lengths, the nominal arc length was used to calculate strain. Gap change is the percentage change from the initial to final spacing between the solid blocks.StraightSegmentCurved SegmentMaterialStrainGap ChangeStrainABS (Planar Print) ABS (Vertical Print) Acrylic<0.1% 1.5% <0.1%61% 64% 59%0.9% 3.1% 2.0%By integrating the transient heating with a compliant mechanism, the local thinning of theflexing regions will heat through more quickly increasing the contrast in stiffness between the thick and thin regions.Spatial Variation in Local StiffnessGiven that the printed material has fairly uniform material properties, variation in localstiffness is achieved by carefully designing the geometry. The key to this approach is designing geometry that is thin and can bend significantly without fracture.

The key to this approach is designing geometry that is thin and can bend significantly without fracture. Often the effects that are desired require nonlinear deformation theory, and these effects are most aptly designed using compliant mechanisms.A mechanism is a device that transforms or transmits motion or energy; a compliant COMPLIANT MECHANISMS mechanism is a mechanism that achieves some or all of its motion through the deflection of flexible members (Howell, 2001). There are various techniques that have been proposed for the design of compliant mechanisms, including topology optimization (Sigmund, 1997), building block (Gallego and Herder, 2009)), and the pseudo-rigid-body model (PRBM) (Howell, 2001; Feng, 2010; Su, 2009). For the work done here, we used the PRBM approach because it lends itself to intuitive design approaches by way of analogy to existing rigid-body mechanisms.

For the work done here, we used the PRBM approach because it lends itself to intuitive design approaches by way of analogy to existing rigid-body mechanisms. This is particularly advantageous when an existing mechanism has the motion desired, the motion of a scissor jack or a collapsible ironing-board.The PRBM approach identifies compliant mechanism topologies that are physically analogous to mechanical linkages. For example, the large deflection motion of a cantilever beam is well-approximated (<0.5% relative error) by placing a pin joint 1/6 of the beam length fromthe fixed end; the stiffness of the beam is simulated by placing a torsional spring at the pin joint. In rigid-body mechanisms, particular emphasis is placed on mechanisms with one or few degrees of freedom (i.e. controllable inputs).

In rigid-body mechanisms, particular emphasis is placed on mechanisms with one or few degrees of freedom (i.e. controllable inputs). The reason is that motion prediction is simplest with only one (or few) input variables.The notion of degree-of-freedom is only approximately valid for compliant mechanisms, in the sense that because of the potential energy associated with the flexible parts, the motion of a compliant mechanism is predictable (if occasionally unstable) even when the compliant mechanism has far fewer inputs than its rigid-body analog. The predictability of the compliant mechanism is based on the fact that its configuration minimizes the potential energy of the system consistent with the work done by the applied loads acting on it.

The predictability of the compliant mechanism is based on the fact that its configuration minimizes the potential energy of the system consistent with the work done by the applied loads acting on it. For example, a paper clip has a known equilibrium configuration when it is unloaded (in the box), when it is loaded (by placing a sheaf of paper in it) it deflects to the position that minimizes its deformation energy consistent the loads the paper places upon it. The scope of designs available for compliant mechanisms is vast and includes spring design, origami, and many micro-electromechanical systems (MEMS) (Howell, 2001).Figure 2 Illustration of a small length flexural pivot.

In the Pseudo Rigid BodyModel (PRBM), it is approximated by a pin joint placed at the center of the thin section.A commonly used compliant ‘joint’ is the small-length flexural pivot, in which a short,SMALL-LENGTH FLEXURAL PIVOT thin segment connects two larger thicker and more rigid segments as illustrated in Figure 1. When bent, the small length flexural pivot acts like a pin joint at the beam center with a torsional stiffness of EI/L, where E is Young’s Modulus, I is the 2nd moment of area, and L is the length of the compliant segment. The advantage of this joint is its simplicity, the disadvantage its weakness, i.e.

The advantage of this joint is its simplicity, the disadvantage its weakness, i.e. it has neither significant bending stiffness nor strength and cannot be expected to resist significant loads.As a demonstration of the possibilities of deformation post-processing using spatialCOMPLEX JOINT variations in local stiffness, we designed a joint which could serve the same function as a small- length flexural pivot, but that had significantly more stiffness and strength. In order to achieve this, we wanted a joint that was larger and capable of accommodating larger deformations. The compliant joint that we designed is a highly redundant parallel mechanism, similar in function to the mechanism used in a scissor jack (for lifting cars) or a pantograph mechanism. Thesemechanisms use repeated parallelograms, and changes in their height and extent are accomplished by changing (shearing) the interior angles of the parallelograms.

Thesemechanisms use repeated parallelograms, and changes in their height and extent are accomplished by changing (shearing) the interior angles of the parallelograms. The joint is designed to accomplish a 90-degree bend with a high degree of strength and repeatability. In the bending of a uniform beam, half of the beam is in tension, half is in compression. In our joint design, we removed entirely the compression half of the beam by putting a 90 degree notch, this allowed for precise bending of the joint because there is a form-controlled mechanical stop. The parallelogram mechanism was placed on the tension side of the joint and allowed for the necessary elongation in that part of the joint.Joint DemonstrationsBending the small length flexural pivots can be as simple as locally heating the materialat the bends. The local heating provides the local variation in material properties that is required. Figure 1 shows a series of parts in their progressive bending states.

The local heating provides the local variation in material properties that is required. Figure 1 shows a series of parts in their progressive bending states. These bends were accomplished by locally heating the creases with a pocket butane torch. Several designs incorporated geometric features to increase the accuracy of the bent part. This approach is simple, but it relies on operator skill for accuracy and parts can easily be damaged by overheating or inaccurate bends. It is also difficult to apply if multiple joints need to be bent simultaneously. If heated through, the joint could be stretched such that the part is permanently deformed.a)b) Figure 3 Deformation post processing failures. a) The part heated until all the sectionswere soft. b) The local joint was heated through until it stretched nearly to failure.Figure 1shows a block that was placed in an oven and heated slowly with and without anapplied force.

b) The local joint was heated through until it stretched nearly to failure.Figure 1shows a block that was placed in an oven and heated slowly with and without anapplied force. The applied force (Figure 1(b)) straightened the joints, but created a tension deformation while heating without an applied load eventually deformed all parts of the model as seen in Figure 1(a).Figure 4 Process of folding a cube from an initial flat state into a full cube illustrating avariety of joint designs. a) As-printed state b) Close up of a bending segment c) torch used to locally heat the flexible segments d-i) documentation of the bend process.Much better performance was obtained by using the complex joints illustrated in Figure2. These joints were placed in a furnace preheated to 200 C and left for 15-20 s. They were removed from the furnace and bent into their final shape and held until they cooled. A water spray was used to speed the cooling process.

They were removed from the furnace and bent into their final shape and held until they cooled. A water spray was used to speed the cooling process. These thin elements of these joints heat more quickly than the bulk material even when the entire object is placed in a hot furnace. Thus the joint deformed readily while the bulk material remained rigid and did not deform. Joint strength could be increased further by solvent welding the flat faces together after deformation.Figure 5 Three compliant joints designed for accurate 90 degree bends. The thinsegments bend readily while the rest of the part remained rigid.3D AntennaNassar and Weller (2013) demonstrated the printing of a 3D antenna using a combinationof SLA structure and direct write conductors. They showed that the antenna gain decreased 2.5 dB relative to a printed circuit board antenna.

They showed that the antenna gain decreased 2.5 dB relative to a printed circuit board antenna. While they expected that reliability might be increased by direct printing, the printing process is relatively difficult and the conductivity of the direct write conductors is reduced relative to bulk metals.We applied deformation post-processing to make the antenna directly from a flat FDM component as illustrated in Figure 4(a). A copper tape was applied to both sides of the FDM- printed surface Figure 4(b). The tape was then cut using a CNC router and the excess removed (Figure 4(c-f)). The process was repeated for the other side and then the FDM part was deformed into its final shape. The deformation process was facilitated by the addition of open regions where the copper crosses the 90° bends.

The deformation process was facilitated by the addition of open regions where the copper crosses the 90° bends. This dramatically reduces the strain required in the copper to perform the bends so that the bends were easily completed without deformation of the copper pattern.Figure 6 Process of making a 3D antenna by deformation post processing. a) Original FDM component b) Part covered by copper tape c) Conductor pattern cut into copper tape d) Part after removing excess copper g) Illustration of copper bent over material void h,i) Final antenna after patterning both copper elements and deforming to final shape.Benefits of Deformation Post ProcessingThe two primary benefits of deformation post-processing of AM components are reducedprinting cost and easier incorporation of multiple materials. Table 2 compares the predicted build times and material usage of printing in the flat condition compared to the final shape for both the (Figure 1) cube and the antenna (Figure 4).

Table 2 compares the predicted build times and material usage of printing in the flat condition compared to the final shape for both the (Figure 1) cube and the antenna (Figure 4). In this case, the difference in materials usage is not very significant because the open geometry minimizes the support material required. In the case of the antenna, the flat model actually uses more support material due to the need to support a wider base.

In the case of the antenna, the flat model actually uses more support material due to the need to support a wider base. However, the printing time is reduced in both cases (substantially in the case of the cube) due to the faster printing in X-Y axes compared to the Z axis.Table 2 Comparison of printing time and costs for flat versus final goemetry.SupportMaterial(cm3)Build Material (cm3)Printin g Time (min)FinalGeometry10.6099.814146CubeFlatGeometry7.46510.26679Change FinalGeometry30%0.664+5%1.07846%12AntennaFlat Geometry1.2131.26511Change+83%+17%8%The 3D antenna demonstrated an easy technique for creating an antenna without the need for an expensive multi-material deposition system. While the specific method of depositing the copper (cover and trim) is not fully automated, there are many other alternative methods to create a circuit on a printed part.

While the specific method of depositing the copper (cover and trim) is not fully automated, there are many other alternative methods to create a circuit on a printed part. The antenna could have been fabricated by printing directly on the part with a conductive ink using a variety of commercial systems or by utilizing custom printed circuits board processes to pattern and etch the copper. The tight corners and orthogonal printing surfaces of the 3D antenna create a substantial manufacturing challenge. However, when printing in the flat position, the printing process is dramatically simplified. The subsequent bending is readily accomplished with simple tools.ConclusionsWe have proposed designing parts for post printing deformation as a path to reducingprinting time, support material requirements, and simplify the incorporation of multiple materials. This approach has been demonstrated in the fabrication of several simple cubes and in the fabrication of a 3D antenna.

This approach has been demonstrated in the fabrication of several simple cubes and in the fabrication of a 3D antenna. These parts showed a consistent reduction of printing time relative to parts printed in their final shape even though these objects were open structures that required relatively little support material. This work could be extended to making complex shapes with continuous curvature variation.ReferencesDollar, A.M. & Howe, R.D. 2006, "A robust compliant grasper via shape depositionmanufacturing", IEEE/ASME Transactions on Mechatronics, vol. 11, no. 2, pp. 154-161.Feng, Z., Yu, Y., and Wang, W., 2010, “Modeling of large-deflection links for compliantmechanisms” Frontiers of Mechanical Engineering in China. 5 (3): 294-301.Flores, N., Cassabona, J., Mendelson, M., Noorani, R. & Ioi, T.

monstrating Paraflow: Interactive fluid dynamics simulation with real-timevisualization for augmented resin 3D printingGABRIEL LIPKOWITZ, Stanford University, USAJOSEPH M. DESIMONE, Stanford University, USAAbstractWhile resin 3D printers are seeing growing adoption in both manufacturing and personal fabrication settings, detecting print failures in real time remains challenging. Object-detection neural networks have shown benefits in a variety of extrusion-based 3D printing methods. Here, we extend such work to resin printing using a physics-informed machine learning data generation pipeline. Our approach leverages our models of the fluid dynamics of the printing process at every slice, in order to synthetically generate a library of print defects. We show such an approach is capable of providing data sufficiently resembling real-world failures to fine-tune a pre-trained custom defect detection neural network that can alert users of failure in real-time.

Finally, to allow novice users to take advantage of our simulation platform, we integrate our tool into an interactive augmented reality interface, which displays simulation predictions to provide guidance on design and machine parameters prior to printing.Forward fluid dynamics simulationAR simulation overlay3D model real-time trackingFig. 1. Summary of our interactive fluid dynamics simulation and visualization tool for augmented resin 3D printing.6631 Introduction and problem formulationResin 3D printing is finding increasing adoption in real-world fabrication settings, not only in manufacturing plants but also among interested hobbyists and everyday makers [38]. However, challenges associated with predicting and correcting failures in such systems preclude their more widespread use, particularly among novice users, hindering the promise of decentralizing manufacturing, personalizing fabrication, and democratizing product design.

While a variety of tools helping users debug print failures are available for extrusion-based printers, both academic [11, 39] and open-source though online platforms such as Instructables and All3DP, comparatively fewer are available for resin 3D printers.In resin-based 3D printing, which has applications in high resolution biomedical models [41], architectural scale models [8], and fashion [35], among other areas, objects are cured through a projection of UV light through a transparent window underneath the resin bath. The separation forces required to subsequently detach the part from the window, or overcome suction forces in the case of continuous liquid interface production (CLIP) [33], are the primary cause of print failure in resin printing. Also known as Stefan adhesion forces, these exert significant stresses on the growing object each layer, and have been studied experimentally in detail by many [24, 12].

Also known as Stefan adhesion forces, these exert significant stresses on the growing object each layer, and have been studied experimentally in detail by many [24, 12]. If left uncontrolled, these forces can cause detachment of the part from the platform (adhesive failure) or delamination of newly cured layers (cohesive failure) [15]), in either case wasting time and material.Recently, we have developed a novel simulation pipeline for modeling resin 3D printing that, at every layer during printing, predicts suction forces to anticipate print failure [19, 18]. Here, we leverage such simulation capabilities in a machine learning pipeline for resin 3D printing defect detection, synthetically generating print failures in virtual worlds in order to train a custom defect detection model.

To enable particularly new users to use our simulation engine, we wrap our tool in an interactive augmented reality interface, displaying simulation predictions in the context of real-world print failures and the machine environment. In sum, our novel contributions are:A physics-informed synthetic data pipeline to detect failures during resin 3D printing; and,An augmented reality interface that overlays such simulation data on the real-world printer context.2 Simulation-guided data generation for real-time printmonitoringWe frame the major class of resin printing failure described above – delamination of the object from the build platform – as an object detection task. Our specific aim is to detect the scenario where an object is, or potentially multiple objects are, no longer attached to the platform, and instead has, or have, delaminated in the vat.

Our specific aim is to detect the scenario where an object is, or potentially multiple objects are, no longer attached to the platform, and instead has, or have, delaminated in the vat. On such object detection tasks, deep learning-based models have achieved state-of-the-art performance in a variety of domains [20], including in the context of defect detection in additive manufacturing [9, 14, 16]. Other researchers have developed neural networks for the purpose of defect detection in additive manufacturing, but these have almost exclusively been in the context of extrusion printing [4]. Moreover, such defect664a!!] a P k [e r u s s e r p0510b!!!!!!!!!c!POTENTIAL PRINT FAILUREGenerate synthetic failure image datasetTransfer learning with pre-trained object detection CNNFig. 2.

2. We predict plausible print failure scenarios with our forward fluid dynamics model (a), illustrated for a range of illustrative mechanical designs (b) using these predictions to generate a synthetic image failure dataset for training a custom object detection model (c).detection models rely on the collection and manual labeling of real-world data [25, 5]; however, this is tedious and cumbersome, and would be unfeasible for a user initiating a print with an unseen design. Recent work in the machine learning community suggests that simulation engines, akin to that which we develop in this work, can also be effectively leveraged for training neural networks for detection purposes, specifically through the use of synthetic data [22]. Here, the use of computer graphics engines for generating high-quality, perfectly labeled image data has been shown to approach real-world data performance on a variety of computer vision tasks [3, 32, 21].

Drawing from this emerging literature, we explore the possibility of using our simulation engine as a synthetic data generation pipeline. Our overall system architecture is summarized in Figure 2, which we explain in more detail below.2.1 Synthetic data generation pipelineWe outline our data pipeline approach in Algorithm 1. In brief, we anticipate likely print failure scenarios with our forward fluid dynamics simulation framework [19, 18], using these scenarios to generate a library of synthetic image failures for training a custom object detection model. This approach is summarized in Figure 3. In addition to simulating print failure objects themselves, for environmental context (including distractors and object aggressors), we obtain a6653D scan of the printer environment, along with exact CAD models of the printer build area and models to-be-printed, to render photorealistic images with segmentation masks.

To improve detection performance on potentially highly variable real-world data – with multiple camera poses set by the user, under potentially variable lighting conditions, and with diverse failure manifestations – we employ domain randomization, which has been shown to be important for bridging the sim-to-real gap and achieving satisfactory performance on real-world detection [31]. Specifically, we systematically vary aspects of the scene such as lighting, camera pose, and material textures. For implementation, we utilize Blender (version 3.3.0), a popular open-source 3D creation software [2], and its versatile Python API. A sample of synthetic images generated by this pipeline, along with real-world counterparts, is shown in Figure 4, for the case of one, two, or three print failures.

A sample of synthetic images generated by this pipeline, along with real-world counterparts, is shown in Figure 4, for the case of one, two, or three print failures. We use this generated synthetic data for transfer learning with a single pass convolutional neural network, Yolov5 [30], pre-trained on 300,000 open-source data images from the COCO dataset [17]; we train Yolov5s specifically. For all training experiments, we use following hyperparameters: batch size of 8, learning rate of 0.01, train-test split ratio of 90-10, momentum of 0.937. Training was performed on NVIDIA Quadro P5000 GPU.Algorithm 1: PRINT FAILURE DETECTION DATA PIPELINE Function Generate Synthetic Data(W, M): /* Printer P, 3D models M */ Initialize W with printer P Initialize S ← {} foreach i = 1, 2, . . . , imax do/* synthetic world W */ /* synthetic images S */ /* data iteration i */ /* 3D model m */ /* slice of suction failure s f */foreach m = 1, 2, . ..

4. Example real (top) and synthetic (bottom) images of print defects, with the latter generated from the user CAD model and a virtual scene of the 3D printer in Blender, and used to fine-tune a custom object detection model from a pre-trained YOLOv5 network. Segmentation masks for one (left), two (middle), or three (right) failures in a single image are shown to the right of their corresponding synthetic images.6672.2 Object detection model system evaluationFor evaluation purposes, we assess the performance of our object detection model in not only detecting suction-related failure, but also in discerning which of potentially multiple prints, and designs, have failed. To that end, we select several designs from the Thingi10k online repository [42] – representing real-world 3D printed designs – to train our model.

To that end, we select several designs from the Thingi10k online repository [42] – representing real-world 3D printed designs – to train our model. To test our model’s ability to discriminate between similar designs, we select variants of a similar design class, mechanical gear components, shown below:14235While we train our object detection model on synthetic data for the superior scalability reasons outlined above, we evaluate the performance of our object tracking model on real-world data. We obtain a sample of suction-related print failures, specifically by artificially perturbing machine parameters and reducing support volumes, increasing the diversity of these real-world data by randomizing their position in the vat. Qualitative results of our defect detection model are shown in Figure 5, and quantitative training metrics in Table 1.

Qualitative results of our defect detection model are shown in Figure 5, and quantitative training metrics in Table 1. As the number of design classes increases, the performance of our model drops, not unexpectedly, as shown in the confusion matrix corresponding to the five test designs shown in Figure 6. This suggests that a larger number of design classes pose greater challenges for accurate object tracking.668Fig. 5. Qualitative indications of the performance of our object detection model fine-tuned purely on N = 500 synthetically-generated images, for the task of detecting one defect (top), two defects (middle), or three defects (bottom). Numbers adjacent to detected design classes represent approximated confidence scores.0 t n i r p1 t n i r p2 t n i r pd e t c d e r Pi3 t n i r p4 t n i r pd n u o r g k c a bprint 0print 1print 2print 3print 4 backgroundTrueFig. 6.

6. Confusion matrix quantifying classification accuracy of the Yolov5 pre-trained object detection model fine tuned with 500 synthetically-generated print defect images with N = 5 design classes, specifically five five 3D models drawn from a subset of Thingi10k online repository (filtered by category "mechanical").Additionally, recent research in the machine learning community suggests that the combination of large synthetic datasets with a smaller sample of real data can lead to the best performance [23].669To that end, we quantify the impact of injecting a small sample of real images of defective prints – specifically, 5 percent the size of our synthetic data – on model performance. Indeed, as shown in Table 1, this leads to significant improvements in model performance.

Indeed, as shown in Table 1, this leads to significant improvements in model performance. We emphasize, however, that such real-world data is tedious to obtain, and potentially unfeasible for entirely new classes of designs, justifying our complementary simulation-guided data generation approach.Finally, we evaluate the importance of various simulation domain randomization factors in our synthetic data generation pipeline on detection accuracy. To do so, we perform ablation studies where we systematically eliminate one domain randomization parameter and assess the impact on model performance. All experiments performed with 1000 synthetic training sample images and pre-trained models fine-tuned for 50 epochs. The results are shown in Table 2. Here it is clear that camera pose is of critical importance, as expected, with model pose and lighting conditions also important to performance.

The results are shown in Table 2. Here it is clear that camera pose is of critical importance, as expected, with model pose and lighting conditions also important to performance. This is in line with recent findings from the machine learning community [31].Our detection model is trained to be orientation-agnostic, detecting objects from a variety of distances and orientations. However, our model does depend on having access to a clear line of sight to the printer vat, which may not be possible if the printer build platform obstructs view of potentially delaminated objects.

However, our model does depend on having access to a clear line of sight to the printer vat, which may not be possible if the printer build platform obstructs view of potentially delaminated objects. To circumvent this, we can modify our print scripts to incrementally lift and lower the build platform during printing to provide access to unobstructed views; the frequency of such interruptions to printing is a balance between speed and resolution of defect detection.3 Simulation visualization: augmented reality interfaceIn addition to alerting the user that a print failure has occurred, we also utilize our simulation framework to provide guidance on correcting such errors in future prints.

To achieve this, we leverage recent advances in industrial augmented reality, which has seen growing popularity as a tool to overlay digital information on physical assets for real-time monitoring and user interaction [28, 29], including with 3D printers of numerous kinds [26, 1, 6] along with other digital fabrication systems [36, 13]. In general, such systems have enabled greater levels of interactivity and intuitiveness in the design and fabrication process compared with existing CAD/CAM workflows [40, 37, 10].To enable similar interaction with simulation results for our system, we leverage the Unity Game Engine (version 2021.3.25f), a widely used cross-platform game development tool that provides a comprehensive suite of features for creating interactive and immersive experiences [34]. We briefly describe our system implementation for our interactive simulation tool is as follows.

We briefly describe our system implementation for our interactive simulation tool is as follows. We load our670simulation results with variable print parameters into the Unity AR scene environment, such that the user can readily toggle through them in real-world visualizations prior to printing, including by interactively zooming in and panning around the to-be-printed model in its accurate physical context.To realistically overlay our numerical simulation data in the physical printer environment, we track the pose (position and orientation) of both the printer platform and vat; specifically, the user aligns a virtual overlay of the outline of the printer component with the real- world object (Figure 7). Tracking of the build platform and vat of the printer is theoretically straightforward, as exact CAD models are available.

Tracking of the build platform and vat of the printer is theoretically straightforward, as exact CAD models are available. However, as these black anodized aluminium printer components themselves do not offer many readily trackable features, we improve tracking by endowing both printer components of with fiducial markers, specifically Aruco markers. Once tracked, we use their positional information to realistically overlay simulation results, organized as children of their respective printer components in the AR scene. This ensures that as the user pans around the printer, the pose of the virtual simulation data remains consistent with the real-world printer. We develop our tracking module with Vuforia (version 10.15.4) developed by PTC Inc., an augmented reality (AR) software development kit (SDK) [27].Fig. 7.

We develop our tracking module with Vuforia (version 10.15.4) developed by PTC Inc., an augmented reality (AR) software development kit (SDK) [27].Fig. 7. Summary of procedure for tracking and pose estimation for AR simulation over- lay, implemented in Unity3D, with tracked printer elements (platform and vat) high- lighted.3.1 AR simulation virtual overlayTwo general cases of example user interactions with our simulation data are shown in Figure 8 and Figure 9. Each visualization illustrates a snapshot of a different simulation view in the AR interface, showing effective tracking whereby the simulation pose remains consistent despite changing user camera orientations in real world. The first case, illustrated in Figure 8, represents the case of when the user encounters a delamination defect, such as one detected in our module outlined in the previous section.

The first case, illustrated in Figure 8, represents the case of when the user encounters a delamination defect, such as one detected in our module outlined in the previous section. Here, the user can align simulations of suction forces during printing with the failure manifestation, helping to explain where defects occurred and at which slices print parameters should be adjusted in future jobs. The second case, illustrated in Figure 9, shows the case of when the user needs to align a new print for injection 3D printing [18], and select injection rates to administer to offset suction.

The second case, illustrated in Figure 9, shows the case of when the user needs to align a new print for injection 3D printing [18], and select injection rates to administer to offset suction. Here, rather than interfacing with an external CAD program, as is typical, the user can view simulation results directly on the platform where printing will occur, potentially aiding in dimensioning and object positioning and rotation prior to printing.3.2 AR simulation application scenariosWe outline several potential applications of our interactive simulation for enhancing resin 3Dprinting workflows:1. Pre-print planning: In the pre-print planning stage, it can be non-intuitive, especially for novice users untrained with CAD systems, to understand how the design and orientation of a 3D printed object impacts its printability.

By overlaying fluid dynamics simulations671abacSim overlay: Q = 2.5 μL / sdcd3D printer tracking modulebSim overlay: Q = 5.0 μL / seefluid pressure [kPa]3D model visualization moduleSim overlay: Q = 7.5 μL / sFig. 8. Example simulation overlays for a print defect, indicated by asterisk, starting with CAD model registration of the printer vat (a-b), and alignment of the virtual slice simulations with the physical model from varying user viewpoints (c-e).Fig. 9. Example simulation visualizations for injection printing, starting with CAD model registration of the printer platform (a), which is used to accurately depict fluidic networks for iCLIP printing aligned to build platform ports (b).

The user can also visualize simulation results of fluid pressures at varying injection rates of 2.5µLs, 5.0µLs (d), and 7.5µLs (e), shown for a variety of camera viewpoints.on a rendering of the to-be-printed object prior to production, obvious design flaws can be highlighted to the user and caught before printing.2. Real-time failure alerts: During the printing process itself, if suction-related print failures do occur, typically users of vat 3D printers only are made aware of the result when they return to the printer. With our object detection tracking system, and with a simple RGB camera without need for calibration, users could be alerted to failures, saving time and material.3. Post-print debugging: Even if suction-related failures are caught, it can be challenging for users, especially novice ones, to understand how and why the printing process failed.

Post-print debugging: Even if suction-related failures are caught, it can be challenging for users, especially novice ones, to understand how and why the printing process failed. By directly overlaying simulation results on the print failure, as shown in Figure 8 - 9, we provide data that can facilitate the debugging process. In doing so, we allow designers to modify their designs, or print orientations and support structures, more intelligibly with the use of simulation data feedback overlaid on the real physical print result.4 Limitations and Future DirectionsHere, we report our ongoing development of an interactive simulation platform for predicting and visualizing suction-related defects in resin 3D printing. We furthermore demonstrate several ways by which users can interact with our simulation predictions in real-time and in the context of a real-world 3D printer.

We furthermore demonstrate several ways by which users can interact with our simulation predictions in real-time and in the context of a real-world 3D printer. While we demonstrate our tool’s object tracking capabilities – which rely upon the automation of simulation-generated synthetic data – can generalize to real-world data, we note several important limitations. First, our object tracking model breaks down when the resemblance of defective models to their corresponding CAD model declines. Moreover, if designs672bear a high degree of similarity to one another, we observe our object detection model to incorrectly track, and confuse, their respective locations. Finally, while we show our system is able to effectively detect delamination failures, failures may also occur whereby the object remains attached to the platform, albeit deformed.

Finally, while we show our system is able to effectively detect delamination failures, failures may also occur whereby the object remains attached to the platform, albeit deformed. In future, we will seek to more closely integrate our machine learning tracking capabilities with our interactive simulation tool for more seamless user experiences and for on-device performance. Currently our system employs iOS-based mobile devices; novel hardware interfaces for augmented and mixed reality applications, including head-mounted displays (HMDs), promise to further enhance the interactivity of our approach.References [1] Hirotaka Asayama, Daisuke Iwai, and Kosuke Sato. “Diminishable visual markers on fabricated projection object for dynamic spatial augmented reality”. In: SIGGRAPH Asia 2015 Emerging Technologies. 2015, pp. 1–2.[2] Blender Foundation. Blender. https://www.blender.org/. Version 3.3.0, Accessed: June 20, 2023. 2023. [3] Steve Borkman et al.

nsification and Distortion in Selective Laser Sintering of PolycarbonateM. Berzins, T. H. C. Childs, K. W. Dalgarno, G. R. Ryder and G. Stein,University of Leeds, LS2 9JT, UKThe creation of a 3D numerical model for real part accuracy simulations is the goal of this work. An improvement of a first attempt to predict densification, which used a classical moving heat source calculation and a viscous sintering model, is reported: it includes a numerical thermal calculation with thermal properties allowed to vary with 'curl' distortion is also reported. The temperature. A study of polycarbonate part development of curl at part edges and with part thickness is followed. Mechanical and thermal reasons are sought as to why the distortion is concentrated at edges.INTRODUCTIONGeometry errors in a selective laser sintered part have two processing causes.

Mechanical and thermal reasons are sought as to why the distortion is concentrated at edges.INTRODUCTIONGeometry errors in a selective laser sintered part have two processing causes. The sintering process allows the sintered envelope to differ from the envelope traversed by the centre of the laser beam; the offset between the envelopes causes linear errors to the extent that the offset is not controlled. Thermal distortions give rise to both linear (shrinkage) and geometric (warping) errors: for polymer parts warping can occur both during building at elevated temperature and later on cooling to room temperature. This paper is about modelling the sintering process and about warping.A previous study of the sintering of polycarbonate parts by a scanning laser beam analysed the creation of thin rectangular plates lying horizontally in the powder bed and with edges parallel and perpendicular to the laser raster direction (1).

Modelling proceeded by replacing the scanning laser spot by an equivalent blade of laser heat and using classical moving heat source theory (1,2) to calculate the temperature / time history of powder in and round the edges of the part envelope. When combined with a viscous powder sintering model (1,3), predictions of part density and size were made. Both density and size errors depended on the laser heat flux (J/mm2) into the surface, given by P/(Us) where P is the laser power, U is the scan speed and s the separation between laser rasters (this has also been found by others (4)). Good agreement with experiment was found when the room temperature conductivity and specific heat of solid polycarbonate were used in the temperature calculation. This is surprising as the polycarbonate was not solid or at room temperature.

This is surprising as the polycarbonate was not solid or at room temperature. This paper reports developments of this work, aimed at establishing what are the key physical properties that must be included in a sintering model eventually to be a practical 3D prediction tool.As for warping, differential thermal contraction can cause uniform curvature of flat plates to develop during post-process cooling of polymer sintered parts to room temperature (5). Another type of curvature develops during the manufacture of flat plates lying horizontally in the powder bed. It is concentrated at the lower edges of the plates and called curl. It does not seem to have been asked why this curvature is confined to edges, nor is it obvious. It has visual similarities with the spalling of thin films from196substrates (6,7) and this may imply that failure of cohesion between a plate and supporting powder is relevant to curl development.

However, spalling differs from curl in that it is not constrained to thin film edges; it is catastrophic: spalling cracks run almost right through an interface. Experimental studies of curling and theory for residual strain and stress development is reported, to try to explain why, in selective laser sintering, the large curvatures that can occur on downward facing nominally flat surfaces are confined to edges.THEORYSintering. Densification depends on temperature and time; temperature depends on conductivity k, specific heat Cp and density; k in tum depends on densification. The flow chart of a computer programme that develops by iteration a consistency between an assumed and calculated sintered density is shown in figure 1. It also allows for radiation heat loss by adjusting this until the heat from the laser equals that flowing into the part plus that lost by radiation.

It also allows for radiation heat loss by adjusting this until the heat from the laser equals that flowing into the part plus that lost by radiation. The temperature calculation, part of the sintered density calculation, can either be the analytical method (1) or a finite difference calculation: a public domain adaptive mesh finite difference package expected to be supported as part of the NAG library has been used (8). A linear 2D version has been implemented, but it has capacity for extension to a non-linear 3D form. Thermal properties varying with temperature (OK) are included ( Cp in J/kg and k in W/moK, simplified from ref. 3):Cp =935 + 2.28Tksolid = 0.0251 + 0.0005Tksin ter = ksolid (p/ Psolid)Curl development. The interfacial stresses set up between a thin layer of Young's modulus E and a rigid substrate, due to a thermal contraction strain E in the thin layer, have been studied by an elastic finite element (fe) calculation.

Two cases have been considered (figure 2): the layer bonded totally to the substrate; and a layer already separated at one end. A cooling melt layer (as in selective laser sintering) is not elastic but visco-elastic. A conceptual model of such a melt layer bonded to an elastic foundation (shown as part of figure 14) has been analysed to determine the evolution of strain with time. It can be shown that the compressive strain Ef in the foundation at time t when the melt has experienced a temperature drop of /j,T since time t = 0 is given by(2)where a is the thermal contraction coefficient of the film, and the moduli are defined in figure 14a. Equation 2 can be solved as an initial value problem, with Ef= 0 at t = O. The fe and strain estimates are used to gain insight into the conditions of curling.EXPERIMENTATIONSintered density experiments.

The fe and strain estimates are used to gain insight into the conditions of curling.EXPERIMENTATIONSintered density experiments. Data reported previously (1) on the dependence ofdensity on P/(Us) has been used to assess the modelling developments.197(1)Curl development experiments. Rectangular plates 90 mm x 26 mm in area and 1, 2, 3, 5, 10, 20 and 50 layers thick have been sintered in a Sinterstation 2000 machine. The plate thicknesses and the profiles of their bottom surfaces have subsequently been measured with a micrometer and stylus profilometer. The nominal layer thickness was 0.125 mm, P = 10 W, U = 1,189 mm/s and s = 0.15 mm. The powder bed was at 154°C. No base was used in these tests: a base is a raft of lightly sintered powder built immediately below a part to anchor its first few layers during building. In further tests, 50 layer thick parts were built to study the effect of bases on curl.

In further tests, 50 layer thick parts were built to study the effect of bases on curl. Bases were built with P = 10 W, U = 1,189 mm/s butwith a range of scan spacings from 0.15 to 0.63 mm.EXPERIMENTAL AND THEORETICAL RESULTSSintered densities. Figure 3 records how sintered densities varied experimentally (.) with laser flux. Figure 3a compares these with predictions from the analytical modelling. The upper and lower dashed lines labelled P (for powder) and S (for solid) are predictions without the iterative density loop of figure 1. P results when room temperature powder· thermal properties are used in the temperature calculation and S when properties of the solid are used. The solid lines all refer to the use of the iterative development: (i), (ii) and (iii) result when thermal properties for 300oK, 600 0 K and 800 0 K are used respectively. (iii) is a good fit with experiment but there is no independent basis for this choice of thermal properties.

(iii) is a good fit with experiment but there is no independent basis for this choice of thermal properties. Figure 3b contains the predictions from 2D finite difference thermal modelling. Curve (i) has constant thermal properties appropriate to 600oK: its agreement with curve (ii) in figure 3a is a check on the numerical method's accuracy. Curve (ii) has thermal properties varying according to equations 1: it closely predicts densities, but slightly overestimates them at low energy fluxes and underestimates them at high fluxes. In all the calculations the inclusion of radiant heat loss had no significant effect on temperature: it can be neglected from sintering considerations for polymers.Curl development. Figure 5 traces the edge profiles of parts 1 to 50 layers thick, where the part co-ordinate system is defined in figure 4.

Figure 5 traces the edge profiles of parts 1 to 50 layers thick, where the part co-ordinate system is defined in figure 4. The 1 layer part has a downward curl but this is an artefact caused by placing a thin (very flexible) upward curling part upside down on a flat plate in order to measure its profile. Figure 6 shows how the derived slopes vary with distance from the edge. Linear slope variation with x indicates a circular profile: all the profiles can be approximated to 2 circular arcs. The 50 layer thick parts showed a uniform curvature far from their edges. Figure 7 shows a measured slope variation with x over the range x > 15 mm. The much lower change of slope with x indicates a much larger radius (~ 6 m) than for x < 15 mm.In figure 6, a region I of steep slope is seen close to the edge (x < 2.5 mm) and a region II of lesser slope for 2.5 < x, mm < 12.5. The region II spreads further the thicker is the part.

The region II spreads further the thicker is the part. The radius as a function of x has been estimated from 1/(d2y/dx2} and is plotted in figure 8. The 2 layer thick part has an edge curled to a radius of 10 mm over a length of 2.5 mm and a further curl developed over x = 2.5 to 4 mm, with a radius increasing to 15 mm. (The 1 layer thick part was found to have only the 2.5 mm length of 10 mm radius.) As extra layers are added, the pre-existing curl is not much changed, but the length of curled region grows and the radius increases, until for x > 5 mm the198radius becomes constant at 55 to 60 mm. Once parts are between 20 and 50 layers thick, and the length of curled region has become 12.5 mm, no further curl occurs. The length e of curled region, considered as having parts eI and en corresponding to regions I and II in figure 6, is further considered in figure 9.

The length e of curled region, considered as having parts eI and en corresponding to regions I and II in figure 6, is further considered in figure 9. The length en stops growing when the part thickness, h, reaches 4 mm; by then, the initial value of enlh of 5 has dropped to 3.The influence of a base on curl profile is shown in figures 10 and 11. Neither the initial radius R1 nor the length e of curled edge is changed by a base (figure 10) but Rn depends on the laser flux used to build the base (figure 11). Figure 11 shows the base has no restraining effect on curl until a flux greater than 0.015 J/mm2 is used. Increasing the flux to 0.03 J/mm2 fully develops the restraint: this flux causes densification in the base to a level of 560 kg/m3 (figure 3).Causes of curling have been searched for.

Increasing the flux to 0.03 J/mm2 fully develops the restraint: this flux causes densification in the base to a level of 560 kg/m3 (figure 3).Causes of curling have been searched for. The edge shape of a single layer part has been estimated from analytical calculations of the temperature / time histories of points near the surface of a powder bed, round the edge of a rectangular envelope scanned by a laser beam. Sintered density contours derived from these histories are shown in figure 12. Curved contours are seen which correspond to a corner of radius of 0.5 to 1 mm, one tenth the value of the initial curl radius of region 1. (Microscopic observation does indeed show such a fine edge radius to exist, but it has no relevance to curl.) Finite element calculations (figure 2) show tensile interfacial stresses to exist for x /h < 0.5 and shear stresses to exist for x/h < 3 (figure 13). They are the same for both the uncracked and cracked interface cases.

They are the same for both the uncracked and cracked interface cases. These xIh values are less than the observed e/h extents. Further, if the stresses are responsible for the curl, the model gives no reason why the curl should be limited to the part edges. Finally figure 14 summarises the viscoelastic residual strain calculations from equation 2. Significant strains only develop once the melt temperature drops to < 190°C. In this rubbery region, the modulus Em1 (figure 14a) was supposed infinite, to create a Voigt model. The melt modulus and viscosity were assumed to vary with temperature as in figure 14b (the viscosity values are extrapolated to lower temperatures from capillary rheometer data in (3), using a WLF transform (9)), and Ef was taken to be 108 Pa (the foundation is at 154°C).

Figure 14c shows a build up of residual strain to reach value from - (3.3 to 6.4) x 10-3 at 155°C, depending on cooling rate from 0.1 to 10 °C/s.The density contours in figure 12 have no allowance made for the shrinkage that the densification implies. The vertical marker indicates a derived shrunk layer thickness of 0.33 mm. The measured thickness of a single layer was 0.61 mm.DISCUSSION AND CONCLUSIONThe thermal sintering calculations have demonstrated that variation of thermal properties with temperature must be taken into account in modelling, but radiation can be ignored. An attempt to predict the thickness of a single layer part (figure 12) underestimates thickness by a factor of 2.

An attempt to predict the thickness of a single layer part (figure 12) underestimates thickness by a factor of 2. At the time of preparing this paper it was being studied numerically whether allowance for densification on thermal properties, concurrently within the temperature calculation, would remove this discrepancy and thus add a further essential factor to be included, as suggested in (4).Part curvature exists on several·scales. Sintering calculations (figure 12) have given radii from 0.5 to 1 mm for an edge extent e < 0.5 mm at the edges of single layers. At the other199between, a curl extreme a general warpage radius of 6 m has been measured (figure 7). radius of 10 mm, not affected by bases, has been measured for e up to mm on single layers. As multiple layers are built, curl extends until e reaches 12.5 mm for a part 4 mm thick; thereafter no further curl develops (figures 8,9).

As multiple layers are built, curl extends until e reaches 12.5 mm for a part 4 mm thick; thereafter no further curl develops (figures 8,9). The radius of curl more than 5mm from an edge is affected by bases, varying from 60 mm with no base to 200 mm (figure 11). We speculate that the two levels of curl have different causes: the 10 mm radius is due to sintering shrinkage while the 60 to 200 mm radii arise from thermal contraction as a melt layer cools back to 154°C. In the latter case, the elastic calculation (figure 13) gives no reason why curl is confined to edges. There must be a cohesion with supporting powder to prevent spread of curl. calculations of figure 13 and the viscoelastic strain estimates of figure suggests a cohesion strength in the range 0.1 to 1 MPa. Figure 14 shows the strain reduces with reducing cooling rate. Perhaps curl is confined to edges because a larger cooling rate at edges causes larger stresses there.

Figure 14 shows the strain reduces with reducing cooling rate. Perhaps curl is confined to edges because a larger cooling rate at edges causes larger stresses there. A full explanation of this, with better viscoelastic modelling, awaits further work.ACKNOWLEDGEMENTOne of us (GRR) wishes to thank the Keyworth Institute of the University of Leeds for aresearch scholarship.1. Childs varying powers, scan speeds and spacings, Proc. Solid 1994, pp 356-363, The University of Texas at Austin 1994. Jaeger J.1. Childs varying powers, scan speeds and spacings, Proc. Solid 1994, pp 356-363, The University of Texas at Austin 1994. Jaeger J.Moving sources of heat and the temperature at sliding contacts, Jni.Soc. New South Wales, 1942,76,203-224. 3. Nelson J. Ind. Eng. Chern. Res., 1993,32,2305-2317. 4. Nelson J. C., Selective laser sintering: a definition of the process and an empirical sintering model, PhD thesis, The University of Austin at Texas, 1993. 5.

NSIFICATION BEHAVIOR OF SLS PROCESSED AI203/Al COMPOSITET. Srinivasa Rao+ ; D.L. Bourell* ; H.L. Marcus** + Regional Engineering College, Trichy-620 015, India * Center for Materials Science and Engineering, The University of Texas at Austin, Austin, TX 78712 ** Institute of Materials Science, The University ofConnecticut, Connecticut 06269ABSTRACTProduction of structurally sound parts by any rapid prototype technique is essential, because fully functional features are necessary where application testing is required. In the present work, a powder blend of A1203/AI (3:1 by weight) was mixed with ammonium dihydrogen phosphate and subjected to selective laser sintering (SLS) using a C02 laser. An attempt has been made to increase the powder bed density by introducing vibration to the part cylinder. These SLS in a hydrogen processed preforms were then subj ected to a secondary heat treatment atmosphere and to hot isostatic pressing.

These SLS in a hydrogen processed preforms were then subj ected to a secondary heat treatment atmosphere and to hot isostatic pressing. Densification behavior of these Al20 3/Al composite preforms is discussed.INTRODUCTIONRapid technological developments are driving increased global competition for compressed design and manufacture of new products. a number of solid freeform fabrication (SFF) techniques have been developed to produce rapid prototype solid objects. to These SFF techniques reduce conventional methods of fabrication of prototype solid parts. However, parts produced by many of these SFF techniques exhibit inferior mechanical properties relative to conventionally produced counterparts. The selective laser sintering (SLS) process, one of the SFF techniques, has commercial feasibility for the production of solid parts using polymeric and wax materials.

The selective laser sintering (SLS) process, one of the SFF techniques, has commercial feasibility for the production of solid parts using polymeric and wax materials. Presently, the SLS process is in the developmental stage relative to production of metallurgically sound and fully functional parts from metals, ceramics and composite materials [1-6]. Functional features in the prototype solid part are essential where applications testing is required. Successful processing of these materials, however, depends on process parameters described elsewhere [7].

Successful processing of these materials, however, depends on process parameters described elsewhere [7]. The aim of the present to study densification behavior of an A1203/AI composite system under the following conditions: a) introduction of vibration to the part cylinder to increase powder bed density before SLS, b) secondary heat treatment of the SLS processed preforms in hydrogen and c) hot isostatic pressing (HIP) of the SLS processed preforms.In recent years,time and cost of production substantially comparedinvestigation isEXPERIMENTAL DETAILSElectronics grade calcined alumina powder and aluminum powder used in the present study were uniformly mixed in the weight ratio of 3:1, respectively. To this powder blend, 20% by374weight added as a binder. The characteristics of these powders are highlighted in the T~ble. After proper blending, the composite mix was subjected to SLS using a C02 laser.

The characteristics of these powders are highlighted in the T~ble. After proper blending, the composite mix was subjected to SLS using a C02 laser. The process parameters maintained during SLS were as follows:ammonium dihydrogen phosphate, NH4H 2P 04 wasIncident Power Beam Diameter=14W =0.5 mmScan Speed Scan Line Spacing = 5 mils (125 !lm)= 6.35 cm/sTABLE: Characteristics of the powders used in thisstudyMaterialAverage particle size (!lm)ShapePurity (%)SourceAlumina15Flake99.55Norton CompanyAluminum20Spherical99.00Johnson Matthey Catalog CompanyAmmonium dihydrogen phosphate* * received in crystalline form and ground to (20!lm size( 20Flake99.95Johnson Matthey Catalog CompanyTo study the effect of powder bed density on the SLS processed preforms, the powder bed density was varied before SLS.

A stainless steel disc weighing 300 gm was placed on the powder bed surface and with an electric engraver on top of it, and the powder bed was vibrated for a period of 0 to 30 seconds to generate a wide range of powder bed density. To remove the binder from the composite system and to study the densification behavior, these SLS processed preforms (with an average relative density of 46%) were subjected to a secondary heat treatment in a hydrogen atmosphere at two different temperatures, 800T and 850°C. Soaking time was varied from 2 to 10 hours and the preforms were furnace cooled. Volumetric shrinkage was calculated by measuring the change in dimensions of the preforms after the heat treatment. An attempt was also made to study the densification behavior of the SLS processed preforms during HIP. Initially, SLS processed preforms were encapsulated in pyrex glass before HIP. The temperature, pressure and duration of HIP were 940±10°C, 10,000 psi and 60 minutes.

Initially, SLS processed preforms were encapsulated in pyrex glass before HIP. The temperature, pressure and duration of HIP were 940±10°C, 10,000 psi and 60 minutes. Density was calculated by dividing the mass of the samples with nominal dimensions of 10.0mm X 10.0mm X 3.0mm (± 0.3mm variation) by their true volume determined from linear dimensions and relative density was calculated by dividing the actual density of the SLS processed preforms by the theoretical density of the composite system.RESULTS AND DISCUSSIONIt tapped or vibrated, there is a change in packing configuration of the powder bed from loose random packing towards close random packing. This results in improvement in the packing density. An attempt was made to study the densification behavior of the SLS processed preforms with the powder bed density. Figure I shows the effect of vibration to the Al20 3/Al + NH4H2P04 composite powder bed.

Figure I shows the effect of vibration to the Al20 3/Al + NH4H2P04 composite powder bed. Curve a represents the density of the powder bed and Curve b represents the SLS processed preform. The relative powder mix in the loose random configuration is 35% and increases up to a maximum of 42% as vibration time is increased. No further improvement in packing density was observed beyondis known that once the powder bed is37525 seconds of vibration time. When the powder bed was subjected to improvement there is a possibility for improving the initial density of the SLS processed preform by as much as 27% by vibrating the powder bed. Figures 2a and 2b are scanning electron micrographs of the SLS processed A1203/A1 composite preform with 46% relative density. demonstrate the wettability of the ammonium dihydrogen phosphate. Figure 2a represents a low region and figure 2b represents a high density region.

demonstrate the wettability of the ammonium dihydrogen phosphate. Figure 2a represents a low region and figure 2b represents a high density region. In both micrographs, it is Al203/Al particles are completely wet and covered by the ammonium dihydrogen phosphate. Despite good wetting, part overall density remained less than 50%.after vibration, indicatesin the relative density of the SLS preform was seen.thatThough ammonium dihydrogen phosphate was used as a binder, like most in the system reduces high-temperature properties of the composite. To remove the composite system, SLS processed preforms were subjected to a heat treatment atmosphere. After the heat treatment, in spite of volumetric shrinkage of the preforms, the overall density of the preforms was not improved, and in some cases, actually decreased. This is attributed to the loss of mass due to the evolution of NH3 and during the heat treatment.

This is attributed to the loss of mass due to the evolution of NH3 and during the heat treatment. Figure 3 shows the effect of the heat treatment temperature and time on the percent volumetric shrinkage. Curve a represents 800°C and Curve b represents 850°C. Volumetric shrinkage increases with increasing temperature and holding time. This increase the volumetric shrinkage is higher at initial stages of holding time and tends toward saturation after 8 hrs holding time. Figure 4 is a scanning electron micrograph (cross section) of the heat treated (850°C for 8 hrs )SLS processed preform. The preform has extended interconnected porosity, and the glassy phase enveloping the A1203/Al particles as shown in figure 2b is removed. However, an earlier study of this binder in SLS [8] indicated that a remains in the system.its presence phase from a hydrogen processedofFigure 5 is a scanning electron micrograph of a HIP SLS processed preform with 76% theoretical density.

In spite of the high temperature and pressure used during HIP, full density could not be achieved for this system, and interconnected porosity is still present.

Generation of ammonia and water vapor in the pores may have contributed to incomplete closure of pores.CONCLUSIONSThe salient features of the present study are: a) improvement in initial density of the processed preforms up to 27% is observed by increasing the powder bed density via vibration of the part cylinder; b) by heat treating the SLS processed preforms, composite system could be removed but with concomitant loss of density ; and c) maximum relative density that could be achieved for this system by HIP is approximately 76%.the glassy phaseACKNOWLEDGMENTSOne of the authors (TSR) would like to thank the Department of Science and Technology, Government of India and Board of Directors of the Regional Engineering College, Trichy-620 015, India and Center for Materials Science and Engineering (The University of Texas at Austin) for sponsoring his sabbatical leave under a BOYSCAST Fellowship.376REFERENCES1. D.L. Bouren, H.L. Marcus, lW.

NSIFICATION OF SELECTIVE LASER SINTERED METAL PARTSBY HOT ISOSTATIC PRESSINGMukesh K. Agarwala, and David L. Bourell, Center for Materials Science and Engineering,Joseph J. Beaman, Department of Mechanical Engineering,The University of Texas at Austin, Austin, TX 78712.ABSTRACTMetal matrix alloy composite parts were made from powders by Selective Laser Sintering (SLS). In this study, partially dense (60%-80%) metal parts made by SLS were densified to full density (>98%) by hot isostatic pressing (HIPping) without any loss of shape. HIPping was done by vacuum sealing SLS samples in glass capsules. HIPping parameters, such as, temperature, pressure, and time, were studied with respect to density, linear shrinkage, and microstructures. Anisotropy in linear shrinkage was correlated to the SLS processing parameters. Densification resulting from HIPping was correlated to microstructures and theoretical HIP densification maps.

Anisotropy in linear shrinkage was correlated to the SLS processing parameters. Densification resulting from HIPping was correlated to microstructures and theoretical HIP densification maps. A detailed analysis of such maps is presented.INTRODUCTIONThe ultimate aim of rapid manufacturing processes is to fabricate three-dimensional fully functional parts directly from metals and ceramic materials, without the use of any intermediate binders or any other materials which may require additional processing steps before or after the rapid prototyping operation. Selective Laser Sintering is one of the few rapid manufacturing processes which possesses the capability of producing such structurally sound parts directly from metals and ceramics [1-4]. Feasibility of producing metal parts directly by SLS has been demonstrated using various metal systems [5-12]. However, SLS parts of metals and ceramics are not fully dense and hence require post processing to obtain near full density.

However, SLS parts of metals and ceramics are not fully dense and hence require post processing to obtain near full density. Typically, infiltration and conventional sintering, such as solid state or liquid phase sintering have been employed so far as post-processing steps for SLS parts [1-4]. These post-processing techniques produce near-full density parts, but still leave behind closed pores and result in anisotropic and unpredictable shrinkage during densification. In this study, SLS metal parts have been post processed to full density by Hot Isostatic Pressing (HIPping).Hot isostatic pressing is a materials processing technique in which high isostatic pressure is applied to a powder part or compact preform at elevated temperatures to produce particle bonding [13-15]. This process usually results in a fully dense body, although partially dense bodies can also be intentionally produced.

This process usually results in a fully dense body, although partially dense bodies can also be intentionally produced. In hot isostatic pressing, the contact between the particles is increased due to deformation by the external forces and the thermal mobility of the atoms. Therefore, with HIPping it is possible to produce a part with a density approaching the theoretical density and with properties of bulk: metal.the exact mechanism driving the densification process can easily be determined during HIPping at a given set of process parameters. Using these equations, densification maps are constructed for various material systems. From such maps, shown later for single phase pre-alloyed bronze powders, it can be easily seen that at low temperatures and short time, yielding is the dominant mechanism by which densification takes place. At high temperatures and long times, power-law creep is the dominant mechanism by which densification takes place.

At high temperatures and long times, power-law creep is the dominant mechanism by which densification takes place. Although such maps exist for various single phase metal and ceramic systems, densification maps for multi-phase systems have rarely been studied.Based on certain constitutive equations [16-18],65Also, such maps do not take into account the effect of phase transformations on the densification process.As the pressing temperatures increases, the amount of pressure necessary for compaction decreases. Typically, HIPping temperatures range from 450°C for aluminum alloys to 1700°C for tungsten. Pressures are usually applied using argon gas as the pressure transmitting medium and ~IPping pr~ssures r~nge from 20MPa to 300MPa: The duration of the pressing cycle is also Important, SInce at hIgh temperatures lengthy pressIng leads to creep.

The longer the pressing is done, the greater the density and the better the properties of the pressing at a given temPerature and pressure. Lengthy pressing also brings about more complete reduction of oxides, softening and recrystallization, which lead to an improvement in the plastic properties of the part. In addition to temperature, pressure, and time, particle characteristics such as particle size also play an important role in densification during HIPping.In this study, HIPping was explored as a post processing route for SLS metal parts. SLS bronze-nickel parts, previously used to study the structural integrity and effect of post processing by conventional liquid phase sintering [10-12], were used for HIPping. SLS parts typically have surface connected and interconnected pores. Therefore, such parts require the surface to be sealed or the part to be encapsulated such that the pressure transmitting medium, argon gas, does not flow into the pores in the part.

Therefore, such parts require the surface to be sealed or the part to be encapsulated such that the pressure transmitting medium, argon gas, does not flow into the pores in the part. Presence of high pressure gas in the interconnected pores will prevent them from "closing" during the HIP processing. To seal such parts, the parts are encapsulated in an envelope such that the "empty space" in the pores is "squeezed" out during HIPping and the pores close to form a fully dense part.The basic requirements for the encapsulation material are that it should be relatively strong, gas tight, inert and plastic under the applied temperature and pressure conditions, compatible with the material to be densified so as to minimize diffusion reactions and easily removable. Several materials and techniques are used for encapsulation of powders and compact preforms for HIP processing [13-15].In this study, the SLS bronze-nickel parts for HIPping were contained in a vacuum-sealed glass capsule.

Glass encapsulation has so far been used only on experimental basis. Commercially, sheet metal containers or ceramic containers are commonly used. Glass encapsulation was used instead of sheet metal containers or ceramic containers because of the ease with which glass can be shaped to take the form of SLS parts irrespective of the complex nature of the part. Unlike sheet metal and ceramics, glass becomes vitreous at high temperatures and hence can flow easily and conform to the shape of the preform or in this study, the SLS part.EXPERIMENTAL PROCEDUREGLASS ENCAPSULATION OF SLS BRONZE-NICKEL PARTSSLS bronze-nickel samples of -70% density were cut from the shoulder section of broken tension test coupons, previously tension tested [lO-12], into small sections of 1.5cm x 1.5cm x O.5cm dimensions. The samples were degassed by annealing them at 450°C in for 2 hours. Commercial borosilicate glass tubes of O.D. 19 mm and LD. 15.5 mm were used as the container material.

The samples were degassed by annealing them at 450°C in for 2 hours. Commercial borosilicate glass tubes of O.D. 19 mm and LD. 15.5 mm were used as the container material. The tube and the sample were thoroughly dried in an oven at IS0°C. One end of the glass tube was sealed by a conventional glass blowing method using an oxy-acetylene torch. The SLS sample was then placed in the tube and a small constriction was made in the tube, -2-4 em above the sample, by the usual glass blowing technique using a torch. Figures lA and IB show schematically the stages involved in vacuum sealed glass encapsulation of the SLS parts. The constriction was as small as possible.66The glass tube with the sample and the constriction was then evacuated using a diffusion pump which was backed up by a mechanical pump, Figure lB. A liquid nitrogen cold-trap was used in the evacuation to ensure efficient evacuation and prevent any contamination of the diffusion pump.

A liquid nitrogen cold-trap was used in the evacuation to ensure efficient evacuation and prevent any contamination of the diffusion pump. The glass tube with the sample was evacuated to a vacuum of -10-5 torr. Once the desired vacuum level was reached, the constricted part of the tube was sealed completely, using the torch, and removed from the remaining part of the tube.HOT ISOSTATIC PRESSING OF SLS BRONZE-NICKEL PARTSHIPping of the vacuum-sealed glass encapsulated SLS bronze-nickel samples was done at three different temperatures of 750°C, 825°C, and 900°C with a pressure of 124MPa (l8ksi) for 1, 2 or 3 hours. Similarly, HIPping of SLS bronze-nickel samples of starting density 70% was done with three different pressures of 69MPa (lOksi), 124MPa (l8ksi), and 180MPa (26ksi) at 825°C for 1, 2 or 3 hours.

Similarly, HIPping of SLS bronze-nickel samples of starting density 70% was done with three different pressures of 69MPa (lOksi), 124MPa (l8ksi), and 180MPa (26ksi) at 825°C for 1, 2 or 3 hours. Sample dimensions and weight were measured before and after HIPping to determine densification.RESULTS AND DISCUSSIONSLS bronze-nickel samples exhibited densification at the various HIPping temperatures, pressures and times used in this study. The extent of densification observed varied with these variables. As shown in Figures 2, density increased with increasing temperatures, pressures and times and approached theoretical density at high values of these variables.Densification of SLS bronze-nickel parts during HIPping was accompanied with phase transformation between bronze and nickel phases, Figures 3 and 4. Bronze, predominantly copper (90 Wt.%), and nickel homogenize at high temperatures by interdiffusion between bronze and nickel to form a homogeneous solid solution.

Bronze, predominantly copper (90 Wt.%), and nickel homogenize at high temperatures by interdiffusion between bronze and nickel to form a homogeneous solid solution. The degree of homogenization depends on temperature and time of sintering or HIPping. As evidenced by (311) peak broadening of copper and nickel, shown in Figures 3, at low temperatures and short times only partial homogenization occurs and complete homogenization occurs at high temperatures and long times. As a result of chemical homogenization, the bronze phase disappears completely leaving behind pores in its place and an expanded solid solution of bronze in nickel. This happens due to a faster diffusion rate of copper into nickel than that of nickel into copper, which results in the Kirkendall effect. Therefore, the pores created at this stage are referred to as Kirkendall porosity.

Therefore, the pores created at this stage are referred to as Kirkendall porosity. Due to the generation of Kirkendall pores and expansion of nickel phase, the samples may undergo growth or swelling instead of shrinkage, as has been observed in SLS bronze-nickel parts during conventional sintering below 1000°C [10-12]. Although, as shown in Figures 3 and 4, chemical homogenozation between bronze and nickel was observed in the SLS bronze-nickel parts during HIPping also, no overall expansion or growth of the SLS parts was observed during HIPping. Parts processed at lower temperatures and pressures, which exhibit low density, show partial homogenization with Kirkendall pores present at the boundaries of homogenized and non homogenized phases, Figure 4b.

However, parts processed at high temperatures and pressures, which exhibit nearly full density, also show partial homogenization but absence of any Kirkendall pores, Figure 4c.As a result of chemical homogenization between bronze and nickel, and generation of Kirkendall pores, the densification of the bronze-nickel samples is considerably slower than predicted by theoretical densification maps for single phase pre-alloyed bronze. Figures 5 show the densification of bronze-nickel samples observed in this study relative to theoretical densification curve for pure bronze of 50Jlm particle size.

However, the experimental data for SLS bronze-nickel parts are well to the right of the theoretical curves indicating a much slower densification in the bronze-nickel samples. Therefore, densification is slowed down during HIPping of mixed phases, such as bronze-nickel, which form a solid solution at HIPping temperatures resulting in Kirkendall pores which oppose the on-going densification. Theoretical HIP maps do not account for such phase transformations and their effect on densification during HIPping. Low HIPping temperatures and pressures and short time of HIPping cause nearly complete removal of pores present due to SLS but do not result in removal of Kirkendall pores created due to homogenization during HIpping. Therefore, such processing conditions which do not remove the Kirkendall pores completely result in less than full density.

Therefore, such processing conditions which do not remove the Kirkendall pores completely result in less than full density. However, high temperatures and pressures result in removal of both the pores due to SLS as well as the Kirkendall pores created due to homogenization during HIPping.Net volume changes (shrinkage) of SLS parts during HIPping was in accordance with the observed density changes. However, the linear dimensional changes, .1.L/Lo, were not the same in all three directions, as observed earlier in conventional liquid phase sintering of such parts. Figure 6A shows the shrinkage in parts HIPped at 825°C and 69MPa for varying times. The linear dimensional shrinkage in these parts was anisotropic with least shrinkage in the scan direction and maximum shrinkage in the thickness or build-up direction. Similar shrinkages have been observed in SLS parts when subjected to conventional liquid phase sintering [10-12].

Similar shrinkages have been observed in SLS parts when subjected to conventional liquid phase sintering [10-12]. In either case, maximum shrinkage in thickness direction is observed due to lowest density along that direction due to poor sintering between layers during SLS and minimum shrinkage in the scan direction due higher degree of sintering during SLS along the scan direction. However, the degree of anisotropic shrinkage observed during HIpping is lower than that observed during liquid phase sintering. Comparision of Figures 6A and 6B shows that the difference in shrinkage between thickness and scan direction is lower in HIpped parts. HIpping is done at temperatures well below the liquid phase sintering temperatures, therefore there is little or no liquid phase formed during HIPping.

HIpping is done at temperatures well below the liquid phase sintering temperatures, therefore there is little or no liquid phase formed during HIPping. Also, HIPping is done under isostatic pressures, therefore there is minimal effect of gravity aiding the densification or flow of liquid in the direction of gravity which coincides with the thickness or build-up direction during both the HIPping as well as liquid phase sintering processes. Therefore, the anisotropy observed in HIPping of SLS parts represents a more precise picture of density gradient present in parts due to SLS.CONCLUSIONSA framework for encapsulation of SLS parts for HIPping has been established using glass containers. Glass encapsulation is an easy, flexible, and economical method for encapsulating complex shapes. Densification studies of SLS bronze-nickel parts by HIpping resulted in higher density parts with increasing temperatures, pressures and times.

Densification studies of SLS bronze-nickel parts by HIpping resulted in higher density parts with increasing temperatures, pressures and times. However, the densification of SLS bronze-nickel parts was slower than that predicted by theoretical HIP densification maps due to generation of Kirkendall pores arising from the chemical homogenization between bronze and nickel phases during HIPping. The anisotropy in linear dimensional changes during HIPping is not influenced by liquid phase or effect of gravity, thus giving a precise picture of anisotropy present in parts due to SLS.ACKNOWLEDGMENTSThis study was made possible by a research grant from State of Texas Advanced Research Project, ATP-116: Selective Laser Sintering: Direct Metal Fabrication71.....J NAPower Law Creep (II) Mech. Driven0.950.950.9c 0;:;.; c 8 '" 0.85 ;> .~~0.8Power Law Creep (I) Mech0.9c or;; c 8 '" 0.85 .::: '"~0.8i:: "UYieldingI ~ONZE 0.75 1· !Part.Dia.

M.K.Agarwala, D.L.Bourell, B.Wu and J.J.Beaman, "An Evaluation Behavior of Bronze-Nickel Composites Produced by Selective Reference # 4, pp.193-202. M.K.Agarwala, D.L.Bourell, B.Wu and J.J.Beaman, "Structurally Sound Metal Selective Laser Sintering," The Congress 1994, the Extraction and Processing Division, At the 123rd TMS Annual Meeting, Mar. 1994, San Francisco, CA, pp. 833-851. M.K.Agarwala, D.L.Bourell, J.J.Beaman, B.Wu, J.W.Barlow, "Selective Laser Sintering of a Bronze-Nickel Powder Mixture," IMS lnt. Conf. on Rapid Product Development, Feb. 1994, Stuttgart, P.E.Rice and S.P.Kohler, "Hot Isostatic Pressing of Metal Powders," Metals pp. 419-443, Vol.7, 9th Ed., Am. H.V.Atkinson, B.A.Rickinson, "Hot Isostatic Pressing," Manufacturing Processes and Materials, Adam 1991. "Isostatic Pressing: Technology and Applications," .....".u.........'"" Elsevier Applied Science, London, 1991.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDEPENDENCE of MICROSTRUCTURE and MECHANICALPROPERTIES on HEAT TREAT CYCLES of ELECTRON BEAM MELTEDTi-6Al-4VS. Raghavan, Mui Ling Sharon Nai*, Pan Wang, Wai Jack Sin, Tao Li, Jun Wei**Singapore Institute of Manufacturing Technology (SIMTech), 73 Nanyang Drive, Singapore 637662Corresponding author:* Email: mlnai@simtech.a-star.edu.sg, Phone: +65 6793 8976 ** Email: jwei@simtech.a-star.edu.sg, Phone: +65 6590 3186AbstractThe EBM Ti-6Al-4V alloy has generally superior mechanical properties, owing to finely spaced α−β laths which give a good combination of strength and ductility. The grain structures in the as-printed structures are long columnar which can give rise to anisotropic mechanical properties.

The grain structures in the as-printed structures are long columnar which can give rise to anisotropic mechanical properties. Moreover the non-uniformity in microstructure can also arise from part geometry where the thin features have propensity to form martensite phase. Heat treatment provides a viable solution to modify the microstructure and to tailor to the properties as desired. A wide range of heat treatment experiments were performed, followed by microstructure and tensile property analyses. It was observed that the microstructure and the tensile properties significantly changed depending on the heat treat cycle performed.

It was observed that the microstructure and the tensile properties significantly changed depending on the heat treat cycle performed. Tensile properties of solution treated air- cool plus aged samples yielded globular equiaxed grains with fine α−β lath structure, which were found to be the best among the different heat treated samples and better than ASTM F1472 specifications.IntroductionElectron Beam Melting (EBM) machine developed by Arcam AB, is capable of producing near net shape complex parts and fully dense Ti-6Al-4V with very fine α−β lath microstructure. The superior microstructural features in the EBM printed parts are mainly due to preheat scan carried out to maintain the powder bed at 650oC-700oC, which transforms the as- printed martensitic α’ structure to an annealed fine α−β lath structure [1, 2].

However, non- uniformity in the microstructure and mechanical properties are observed with regard to the part’s height [3], part’s dimensions [4] and part’s geometries [1] due to the different thermal conditions experienced during the build process. Moreover, the high thermal gradients developed during solidification produce long columnar grains of a few mm length along the build direction [1, 5] which can also contribute to anisotropic properties in the printed part [5].Heat treatment provides a viable method to modify and tailor microstructures to alter the mechanical properties for the applications desired. There has been very limited study on post heat treatment of EBM printed Ti-6Al-4V to improve the properties other than the HIP treatments, possibly because of the superior mechanical properties of the EBM printed parts and any standard705heat treatment does not improve the properties significantly as recently published by Charlotte et al. [6].

[6]. On the other hand, selective laser melting printed Ti-6Al-4V parts are completely martensitic and require a compulsory post heat treatment operation to improve the properties [7]. In the current research, a comprehensive set of heat treatment cycles were designed to study the influence of the process on the microstructure and eventually the tensile properties. The best heat treatment practice of EBM Ti-6Al-4V was also identified which gives refined microstructures for superior tensile properties in the alloy.Materials and MethodsThe samples for post heat treatment study were fabricated in Arcam A2X EBM machine using the Ti-6Al-4V Grade 5 powders with powder size of range 45-105 µm and build layer thickness of 50 µm. Rectangular bars (10 x 30 x 100 mm) parallel to build direction (Z) and perpendicular to build direction (XY) were fabricated as schematically shown in Figure 1, with the build direction marked by an arrow.

The bars were later machined to ASTM E8 dimensions for tensile testing under the as-printed and post heat treatment conditions. The heat treatments were carried out in a vacuum furnace (Thermal Technology LLC) which can attain high vacuum levels of 1e-5 torr. For samples that required air cooling, the samples were vacuum sealed in a quartz tube and heat treated in a Nabertherm muffle furnace. For microstructural observation, the samples were ground, fine polished and etched using Kroll’s reagent (2%HF, 4% HNO3 in water). Images were taken using the Olympus GX51 optical microscope for heat treated samples and Zeiss Field Emission Scanning Electron Microscope (FESEM) for as-printed sample, since the α lath morphology is very fine. The tensile tests were also carried out using the Instron 5982 machine with the strain rate of 3.3e-4 s-1.Figure 1.

The tensile tests were also carried out using the Instron 5982 machine with the strain rate of 3.3e-4 s-1.Figure 1. The orientation of rectangular blocks printed by EBM for tensile tests.The heat treat experiments conducted in this study were classified to three categories: a) sub β-transus, b) super β-transus and c) solution heat treated, air cooled followed by ageing heat treatment. The classification is done to mark the difference in the process conditions and also to emphasize the difference in the observed microstructure and mechanical properties. Figure 2 schematically explains the regions for different heat treatments in the phase diagram and Table 1 shows the different heat treatments performed.706Figure 2.

Figure 2 schematically explains the regions for different heat treatments in the phase diagram and Table 1 shows the different heat treatments performed.706Figure 2. Schematic showing different heat treatment regions.Table 1.Heat treatment conditions used.Heat TreatmentSub β-transus Sub β -transus Super β -transus Solution Air cooled and AgedTemperature/time900oC/2h 950oC/2h 1050oC/2h 1030oC/2h (Air cooled) 500oC/8h (Aged)Orientation & no. samples XY-3, Z-3 XY-3, Z-3 XY-3, Z-3 XY-3, Z-3Results and DiscussionThe microstructure of as-printed EBM sample was long columnar prior β grains with very fine α−β lath structure as shown in Figure 3, which was in good agreement to that observed by other researchers [1, 8-10]. The SEM image of the as-printed microstructure (Figure 4) revealed both lamellar and Widmanstätten morphologies with dark α phase and bright β phase.

The SEM image of the as-printed microstructure (Figure 4) revealed both lamellar and Widmanstätten morphologies with dark α phase and bright β phase. The optical images of microstructures for the different heat treatments taken along the build directions are shown in Figures 5-8 for 900oC/2h, 950oC/2h, 1050oC/2h and 1030oC/2h air cooled and aged 500oC/8h, respectively. A distinct difference was observed in the microstructures for the different heat treatments. The sub β-transus heat treatments still displayed columnar grain structures that are similar to the as-printed microstructure. The β-transus temperature is at 995- 1000oC [11] where the α−phase completely dissolves into β-phase. However, in the sub β-transus temperatures, the presence of α-phase restricted the growth of prior β grains.

However, in the sub β-transus temperatures, the presence of α-phase restricted the growth of prior β grains. Coarsening of α laths occurred in the sub β-transus heat treatments due to diffusion of elements and also the precipitation of α phase in the grain boundaries with increasing temperature was observed. In contrast, the grain structures were completely modified from columnar grains to equiaxed grains after the super β-transus heat treatment. This was caused due to the absence of α-phase in the707matrix and the long columnar β grains were free to rearrange into a more equiaxed morphology, which has a lower energy and stable morphology. The α lath width for the super β-transus heat treatment principally depended on the cooling rate from the solutionizing temperature (above β- transus temperature), hence the furnace cooled samples exhibited coarser α laths (Figure 7) as compared to air cooled samples (Figure 8).

The solution air cooled samples were aged at 500oC for 8h to remove any residual stresses and any martensite that may have been formed. Of all the heat treatments, the solution air cooled and aged samples exhibited equiaxed grains with very fine α−β lath structure which are good for mechanical properties. The α lath width is very critical in determining the strength of the alloy because with fine α−β lath structure, the dislocation slip length is reduced which contributes to increase in strength in the alloy.(a)(b)Build DirectionFigure 3. The microstructure of as-printed Ti-6Al-4V showing (a) columnar grains and (b) fine α−β lath structure.Figure 4. FE-SEM image showing white streaks of β and dark α phases.708(a)(b)Build DirectionFigure 5. Sub β-transus heat treatment at 900oC/2h, showing (a) columnar grains and (b) coarsened α−β laths.(a)(b)Build DirectionFigure 6.

Sub β-transus heat treatment at 900oC/2h, showing (a) columnar grains and (b) coarsened α−β laths.(a)(b)Build DirectionFigure 6. Sub β-transus heat treatment at 950oC/2h, showing (a) columnar grains with α precipitation in grain boundaries and (b) coarsened α−β laths.(a)(b)Build DirectionFigure 7. Super β-transus heat treatment at 1050oC/2h, showing (a) fully equiaxed and (b) very coarse α−β lath structure.709(a)(b)Build DirectionFigure 8. Solution air cooled and aged heat treatment showing (a) fully equiaxed and (b) very fine α−β lath structure.Tensile tests were carried out for the samples of all the different heat treatments and for the samples’ orientations XY and Z to the build plane. The tensile results for ultimate tensile strength (UTS), yield strength (YS) and elongation to failure (El %) are presented in Figure 9. The as-printed EBM samples exhibited the best combination of strength and ductility.

The as-printed EBM samples exhibited the best combination of strength and ductility. The solution air cooled and aged samples also displayed good UTS and YS, which is close to that of as-printed sample. It is interesting to note that only the as-printed samples and the solution air cooled, aged samples exhibited strength higher than the ASTM1472 (wrought and annealed) values in both build directions, marked by the line in Figure 9. The reason for the superior tensile strengths of the as-printed samples followed with solution air cooled samples is attributed to the fine α−β structure. It is also noteworthy that 900oC/2h heat treated samples exhibited the best ductility (~19%) among all the samples. With increase in temperature for sub β-transus heat treatments, the α laths coarsened and the precipitation of α phase on grain boundary also increased, which contributed to the drop in strength for 950oC/2h heat treat samples, as seen in Figure 9.

For the super β-transus heat treatment 1050oC/2h, the samples exhibited the worst properties as the microstructures displayed coarse α−β structure and coarse β-grains due to the high heat treat temperature and slow furnace cooling.In Figures 10-12, the fracture surface of the three heat treated samples (900oC/2h, 1050oC/2h and 1030oC/2h air cooled and aged) are presented. The 900oC/2h fracture surface exhibited a complete ductile failure with the presence of microvoids that led to cavitation and dimpled texture throughout the fracture surface, as seen in Figure 10. On the other hand, the super β-transus heat treatment 1050oC/2h exhibited a mixed failure with fracture surface showing dimples and cleavage planes that indicated brittle failure, as seen in Figure 11. A similar mixed fracture surface with dimples and cleavage planes was also observed for the solution air cooled (1030oC/2h) and aged samples, as shown in Figure 12.

A similar mixed fracture surface with dimples and cleavage planes was also observed for the solution air cooled (1030oC/2h) and aged samples, as shown in Figure 12. The 1050oC/2h heat treated and the solution air cooled samples showed an elongation at failure of 12.1-13.9% which was significantly less than 17.8-18.3% elongation exhibited by 900oC/2h heat treated sample. This difference in ductility was substantiated by the observed difference in the fracture surface of the samples.710Figure 9. The (a) ultimate tensile strength (UTS) (b) yield strength (YS) and (c) elongation % of the different heat treated samples in both XY and Z orientations.Figure 10. Fracture surface of sub β-transus , 900oC/2h sample showing ductile failure with dimples and microvoids.711Cleavage planesFigure 11. Fracture surface of solution super β-transus , 1050oC/2h showing predominantly cleavage planes with some dimples, indicating mixed failure.Cleavage planesFigure 12.

Fracture surface of solution super β-transus , 1050oC/2h showing predominantly cleavage planes with some dimples, indicating mixed failure.Cleavage planesFigure 12. Fracture surface of solution air cooled and aged sample showing mixed failure with both dimples and cleavage planes.ConclusionsThe as-printed EBM Ti-6Al-4V exhibited long columnar grains with very fine α−β microstructure that contributed to its superior tensile properties. For the sub β-transus heat treatments, coarsening of α laths and precipitation of α on grain boundaries occurred with increase in temperature which contributed to loss in strength with increasing temperature. In the super β-transus heat treatment, the absence of α phase allowed the grain morphology to completely change from columnar to equiaxed. The α lath size is mainly controlled by the cooling rates and a clear difference was observed between the furnace cooled and the air cooled samples.

The α lath size is mainly controlled by the cooling rates and a clear difference was observed between the furnace cooled and the air cooled samples. The best refinement in microstructures was observed for the solution air cooled and aged samples which also exhibited the best combination of tensile strength among all the heat treated samples. One of the potential post-processing operation in future could be to combine the solutionizing, quench process with HIPping process to obtain a good microstructure with reduced porosity.712AcknowledgmentsThe authors are grateful for the financial support provided by A*STAR IndustrialAdditive Manufacturing Program: Work Package 3 (Electron Beam Melting), grant no. 132 550 4103.References[1] P. Wang, X. Tan, M.L.S. Nai, S.B. Tor, J. Wei, Spatial and geometrical-basedcharacterization of microstructure and microhardness for an electron beam melted Ti–6Al–4V component, Mater. Des. 95 (2016) 287-295.[2] X. Tan, Y. Kok, Y.J. Tan, M.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDEPICTION OF SUPPORT STRUCTURES IN TECHNICAL DRAWINGSS. Lammers*†, T. Koers*†, B. Magyar*†, D. Zimmer*†, T. Lieneke*†Chair of Design and Drive Technology, Paderborn University, 33098 Paderborn, Germany †Direct Manufacturing Research Center (DMRC), Paderborn University, 33098 Paderborn, GermanyAbstractTo ensure uniform documentation of support structure information, a concept is presented that enables a standardized depiction of support structures in technical drawings based on ISO 128-3. To this end, requirements for a uniform depiction are defined and a procedure for drawing entry is presented. The drawing entry should contain all production-relevant support structure information.

To this end, requirements for a uniform depiction are defined and a procedure for drawing entry is presented. The drawing entry should contain all production-relevant support structure information. The standardized documentation of support structure information in technical drawings is intended to ensure a simple, clear and safe exchange of information between business units or different companies along the value chain. As a result a possible drawing entry of support structures was developed. To distinguish between different support structure types, a standardized depiction of geometrical information in a specification field is shown. The specification field gives a detailed description of the support structure type, the geometry as well as the connection to the part and the building platform. Also uncommon support types like lattice structures or CAD based support structures can be implemented.

Also uncommon support types like lattice structures or CAD based support structures can be implemented. To ensure the usability the depiction is editable and extendable.IntroductionLayer-by-layer and tool-free additive manufacturing processes allow the production of almost all conceivable component geometries. Nevertheless process-specific restrictions must be considered. For example, thermal residual stresses occur in components during metal powder-based laser beam melting, which can result in component distortion and dimensional deviations [1,2]. This can lead to an abort of the build job. In addition, the production of overhanging component areas is only possible with support structures. The support structures ensure better heat dissipation and provide a physical, stable anchorage between the building platform and the component overhang. However, this results in disadvantages, as the support structures lead to increased material consumption and require cost-intensive removal.

However, this results in disadvantages, as the support structures lead to increased material consumption and require cost-intensive removal. In addition, the support structures can have a negative impact on the component surface roughness. Consequently, it is important to use application-adapted support structures for a successful and efficient manufacturing process [3]. It is important to design the support structures in a way that the component requirements, such as costs, surface quality and dimensional accuracy, are met and, at the same time, stability and heat dissipation are ensured during additive manufacturing. Not all requirements can be implemented at the same time. For example, particularly stable support structures lead to robust production, but can only be removed with increased effort. In order to design the support structures to suit the application, there are a large number of different support structures, each of which can be varied via geometric parameters.

In order to design the support structures to suit the application, there are a large number of different support structures, each of which can be varied via geometric parameters. This manufacturing-relevant support structure information has not yet been recorded in a uniform and generally applicable manner. Although it is possible to determine the support structure information from the job preparation program for additive manufacturing, these data takes up a large storage volume and requires expensive software programs to read out the required support structure parameters. Especially, if the used support structure has an influence on the post-processing of the part [4, 5], it is difficult to transfer this information. This is where technical drawings (both in written and digital form) have become established in recent decades. In addition to the component shape and dimensions, materials, manufacturing processes and permissible deviations are also specified in technical drawings.

In addition to the component shape and dimensions, materials, manufacturing processes and permissible deviations are also specified in technical drawings. Thus, they provide a complete, unambiguous, and easily understandable description of components, which are used in different divisions and companies for communication and information transfer [6,7]. An unambiguous and generally understandable depiction is provided by binding rules during drawing creation. These rules are listed in internationally valid standards [ISO 128-2, ISO 128-3].676Objective and procedureIn order to ensure uniform documentation of support structure information, a concept is presented in Figure 1 which enables a standardized form of depiction of support structures in technical drawings based on ISO 128-3. For this purpose, laser beam melting and basic support structure information are explained in the state of the art.

For this purpose, laser beam melting and basic support structure information are explained in the state of the art. Requirements for a uniform support structure depiction are defined and a procedure for drawing entry is presented. The drawing entry should contain all production-relevant support structure information. Not only thin- walled support structures, as used in common support preparation programs, are to be considered, but also special support structure shapes and future support structures should be integrable. The standardized documentation of support structure information in technical drawings should ensure the exchange of information between business units or different companies along the value chain. Integration into software tools (CAD, PDM programs) has to be taken into account by generation of a low data volume.

Integration into software tools (CAD, PDM programs) has to be taken into account by generation of a low data volume. The training effort of the drawing-creating employees should be kept low by applying well known principles of technical depiction.Figure 1: Flow diagram of paper conceptState of the art: Procedure for additive manufacturingFor manufacturing by laser beam melting, a virtual 3D CAD model of the component has to be available. These 3D data can be generated using CAD software or by a 3D scan of an existing object. Likewise, numerical methods, such as topology optimization, enable the generation of a component structure that is suitable for the application loads [8,9,10,11].

Likewise, numerical methods, such as topology optimization, enable the generation of a component structure that is suitable for the application loads [8,9,10,11]. During 3D data generation, design guidelines should already be taken into account for a manufacturing- or post-processing-oriented design of additively manufactured components in order to keep process-related, negative restrictions and the support structure or finishing effort low [4,9,10,12,13,14,15]. For further data preparation, the CAD model is converted into a data format that can be processed by the data processing software of the additive manufacturing system. In most cases, the STL format is used, which triangulates the CAD solid model and converts it into a surface model consisting of triangular facets. If necessary, surface defects caused by the conversion must be repaired in the STL file [14,16].

If necessary, surface defects caused by the conversion must be repaired in the STL file [14,16]. In addition to the STL format, there are other data formats, which, however, are hardly used in the field of laser beam melting. The AMF and 3MF formats should be mentioned here. These formats allow triangulation with curved triangles as well as the storage of additional component information, such as the color or the material [17,18]. The triangulated component model is then virtually oriented in the building chamber of the manufacturing machine. The orientation has a serious effect on the part quality [15,17]. Thus, with a suitable component orientation, the surface quality of the component or the support structure requirements can be optimized. After the part orientation, the support structures are usually generated. This is done with the data preparation software. For this purpose, overhanging areas are automatically supported with support structures.

This is done with the data preparation software. For this purpose, overhanging areas are automatically supported with support structures. Often, support structures are applied to surfaces with an angle of 45° to the build platform. Recent investigations and optimized support parameters [1,18] show that smaller angles are also possible. Ultimately, the limit value for the angle to be supported depends on the material, the machine and the manufacturing parameters used. For example, to improve the detachability of the support structures or to minimize their material costs, various settings can be made to adapt the support structure geometry677to the part or process requirements. Additional solid support structures to improve heat dissipation and warpage can be added if they have not been added in the CAD program during component design.

Additional solid support structures to improve heat dissipation and warpage can be added if they have not been added in the CAD program during component design. After the support structures have been generated, the part can be placed on the virtual building platform with other parts and the final data preparation step of slicing can be performed. During slicing, the virtually created build job is converted into individual layers in which the manufacturing parameters, such as the layer thickness, the component layer contour or the exposure strategy are specified [15,18]. After that, the data can be transferred to the manufacturing system and manufacturing can be initiated. In laser beam melting, a metal powder is applied layer by layer, melted and bonded to the previous layer. A laser scanner unit ensures that powder is melted only locally where the desired components have been placed in the virtual build space. The surrounding, unmelted powder can be reused after a sieving process.

The surrounding, unmelted powder can be reused after a sieving process. The process takes place under an inert gas atmosphere. After all layers have been generated, the components generated in this way can be freed from powder, detached from the build platform and transferred for further post-processing, such as support structure removal or surface treatment.State of the Art: Support structure design parametersThere are many different types of support structures. In the following, the widely used block support and the associated adjustment parameters are shown (Figure 2). It can be seen that there are already a large number of adjustable parameters for this type of support, which have an effect on the process stability, the component quality and the post-processing effort and thus also on the costs. Figure 2 lists only the most important support structure parameters. Depending on the software used, additional settings can be made.

Figure 2 lists only the most important support structure parameters. Depending on the software used, additional settings can be made. The parameters can either be defined globally for the entire build job or an individual support structure can be provided locally for specific component areas. This way, the tooth geometry, perforation or fragmentation can be set, and material consumption and removability can be simplified by a more filigree support structure with plenty of free space between the melted areas.Figure 2: Block support parameters [LLZ21a]678In return, however, thermal conductivity can be impaired, possibly jeopardizing robust production.

Here it is necessary to investigate the different support settings in a structured way and to quantify their advantages and disadvantages and make them comparable, which has only been done in some scientific studies [1,19]Requirements for support structure depictionAs can be seen from the previous section, there are many parameters which can be varied regarding the support structures, and which must be documented for an information transfer. In order to determine the requirements for a suitable support structure depiction, surveys were carried out in scientific and industrial environments. Primarily users of laser beam melting were interviewed, but also participants in the field of standardized, technical representation. During the survey, the following properties were defined as elementary. The support structure representation should be as intuitively understandable, simple, unambiguous, and generally applicable.

During the survey, the following properties were defined as elementary. The support structure representation should be as intuitively understandable, simple, unambiguous, and generally applicable. It should be easy to integrate into the existing standardized drawings and must not contradict existing standards or encourage confusion with other drawing entries. At the same time, it should be possible to represent different types of support structures and to record the support structure parameters that have been set. The representation should be integrable into existing software tools for technical drawing and thus, if necessary, can be added subsequently to already existing technical drawings. Support structures often have a much more filigree and complex structure than solid components. This complex, often lattice-shaped structure requires an enormous data volume in the CAD model. Therefore, a simplified representation of the support structures is useful.

This complex, often lattice-shaped structure requires an enormous data volume in the CAD model. Therefore, a simplified representation of the support structures is useful. This simplified representation is not intended to show the exact shapes and dimensions of the support structures, but to highlight those areas where support structures are found. An exact definition of the support structures can be done by symbols and indication of the parameters and dimensions in supplementary text fields. In this way, the memory and storage requirement can be reduced, and the relevant geometric data can be read directly.Depiction of support structures in technical drawingsAreas of different materials or special component areas are symbolized in technical drawings by different types of lines and area hatching [ISO 128-2, ISO 128-3]. This approach is also applicable to support structure areas.

This approach is also applicable to support structure areas. Thus, support structure areas in technical drawings can be delimited by different contour and filling of component areas.Depiction form of the contoursSimple, wide solid lines were selected for the boundary of the support structure areas. This type of line is also used to represent visible edges and outlines of components and enables a recognizable boundary of the support structure areas as well as clear dimensioning. Dotted lines or zigzag lines, on the other hand, are less clearly dimensionable and are used for the technical representation of other areas, so that there would be a risk of confusion. A combination of several line types should be avoided to enable simple and fast contouring. The contour color is black, which allows integration into existing technical drawings.

A combination of several line types should be avoided to enable simple and fast contouring. The contour color is black, which allows integration into existing technical drawings. A color accentuation in a CAD environment is also conceivable.Depiction form of the area fillingThe filling of the support structure areas ensures differentiation from the actual component. For this purpose, a hatching is used which allows differentiation from existing hatchings in technical drawings [ISO 128‑3] and at the same time symbolizes a simplified depiction of the support structures. After testing various fillings for support structure areas, a cross-hatching was identified as suitable. Although this has some similarity to the standardized hatching used to represent plastics, the use of a significantly larger 65° angle of the cross- hatching allows a clear distinction to be made.

Although this has some similarity to the standardized hatching used to represent plastics, the use of a significantly larger 65° angle of the cross- hatching allows a clear distinction to be made. In addition, cross-hatching resembles a stylized, perforated support structure and thus increases the recognition value. The hatching lines are black with a clear white area between the lines to allow easy integration into existing drawings also by hand. An exemplary drawing entry is shown in Figure 3 a.679Alternative simplified depictionA simplified depiction can be used if support structure areas run directly from the building platform to the component surface. For this purpose, the entire support structure areas are not displayed. Only the supported surface is marked using a zigzag line delimited by dimension arrows.

For this purpose, the entire support structure areas are not displayed. Only the supported surface is marked using a zigzag line delimited by dimension arrows. The angle of the zigzags is 65° (see Figure 3 b.Figure 3: a) Depiction form of the support area filling b) simplified depiction of the support area fillingSpecification of support structure propertiesTo document the exact design form of the support structures and to specify the support structure parameters used, a representation scheme is presented below. Similar to the tolerance specification, a reference symbol is used to mark the corresponding area. The tangible support structure parameters are listed in a specification field. The reference symbol is a rhombus (see Figure 4 a)). The assignment of a support structure area bounded by its contour line is done by connecting the rhombus with a thin reference arrow.

The reference symbol is a rhombus (see Figure 4 a)). The assignment of a support structure area bounded by its contour line is done by connecting the rhombus with a thin reference arrow. The arrow points perpendicularly to the edge of the area to be described, as is usual for other technical drawing entries. Alternatively, the hint arrow can end in the middle of the support structure area. For this purpose, a point should be provided at the end of the line instead of an arrowhead. The arrow is again connected to the rhombus symbol vertically at the left or right tip of the rhombus. (see Figure 4 b) and c))a)b)x x x v vc)xxxxvvd1)d2)d3)Specification fieldSpecification fieldSpecification fieldFigure 4: a) Rhombus reference symbol b) Connection to support area c) Connection to Rhombus symbol d) Possible positioning of specification field (v allowed, x not allowed)The support structure parameters are specified in a rectangular specification field.

This field can directly follow the rhombus symbol associated with the support structure (see Figure 4 d1). If there is not enough space, similar to the roughness specification, the rhombus symbol associated with the support structure can be given a number and the specification field can be placed in a free area of the drawing sheet, referencing the number (see Figure 4 d2). If all support structures have the same properties, it is sufficient to specify a single specification680field. This field does not need to be connected to a support structure area and can be specified above the title block. (see Figure 4 d3). The specification field is outlined with a narrow solid line. Depending on the support structure, it can be divided into several lines with additional information.

(see Figure 4 d3). The specification field is outlined with a narrow solid line. Depending on the support structure, it can be divided into several lines with additional information. Figure 5 shows as an example the specification field for a block support.Figure 5: Specification field for block supportBy using the symbols shown in Table 1, information on the support structure type or geometric features such as the tooth geometry can be symbolized. By adding dimensional information, support structure properties can be characterized. These dimensions must be given in millimeters. If several properties are dimensioned, the dimensions are usually separated by a small "x". If they form a group, as shown in Figure 5, areas 5, 6 and 7, the dimensions are separated by a slash. The first line of the specification field starts with a symbol according to Table 1 a) and b), which identifies the support structure type used.

The first line of the specification field starts with a symbol according to Table 1 a) and b), which identifies the support structure type used. The further structure of the specification field depends on the support structure type.Wall-based support structuresThe support structure symbol in the left column of Table 1 a) indicates that the support structure area does not have a contour wall or the contour wall has different attributes than the hatch area. If the contour wall has the same properties as the support structure in the hatch area, then the symbol version shown in the second column of Table 1 a) should be used. If the contour wall of a support structure area differs in its nature from the hatch walls. This can be indicated by an additional line in the specification field. For this purpose, an unfilled square is entered as a symbol for the specified contour geometry. The specification is done equivalent to the hatch walls.

For this purpose, an unfilled square is entered as a symbol for the specified contour geometry. The specification is done equivalent to the hatch walls. Also for the contour further fields for the thickness of the walls, the fragmentation, the perforation or the tooth geometry can be provided. An example of this is given in Figure 7. The symbol for the support structure type is followed by the specification of the support structure dimension. For wall based block support the relevant dimension is the wall thickness (see Figure 2, Table 1).Table 1: Symbols for support structure typesa) Wall-based support structuresb) Volume-based support structuresBlockMassiveWebContourUsed symbol depends on the existence of contour walls. Left column without contour. Right column with contour.PillarConeLineTreePointLatticeAngleOthers, CAD basedThis is indicated by a preceding small "bw" followed by the dimension in mm.

Right column with contour.PillarConeLineTreePointLatticeAngleOthers, CAD basedThis is indicated by a preceding small "bw" followed by the dimension in mm. By using the symbols listed in Table 1 further support types can be documented and quantified by implementing dimension information. The sequence must be observed. To describe the tooth geometry of the wall-based support structures, the required681information is added in a new line within the specification field. This line is shown in Figure 6a. The double triangles point up and down symbolize the connection to the component and the connection to the building platform. The teeth thus have the same dimensions. If only one pair of triangles is shown, only the tooth geometry to the build plate (downwards) or to the component (upwards) is specified. If no explicit information on tooth geometry is given, the support structure walls extend unchanged to the component or building platform.

If no explicit information on tooth geometry is given, the support structure walls extend unchanged to the component or building platform. The double triangle symbol is followed by the indication of the wall thickness of the teeth. This is indicated by a preceding "bw" and can differ from the main support attributes. The dimensions of the tooth geometry follows in the next section. The base length of the teeth lb, the top length of the teeth lt, the tooth height hz and the interval distance Δb between two teeth are displayed (see also Figure 2).Figure 6: a) Tooth geometry specification field b) Possible tooth synchronicitiesThe support structure synchronicity can be recorded in the specification field. “A” means Asynchronous and that the teeth are cutting each other in a random way. Synchronicity is marked by an “S”.

“A” means Asynchronous and that the teeth are cutting each other in a random way. Synchronicity is marked by an “S”. There are two different synchronicity variants “S1” means that they cut each other in the center of the tooth top and “S2” means that they cut each other in the middle of two adjacent tooth tops. The three variants are shown in Figure 6 b).Structure of specification fieldsSome examples for a depiction in drawings of other wall-based support structures and volume-based support structures are given in Table 2. The general structure of the specification field is the same but there are different attributes that are added into the specification field. Moreover, the introducing symbol from Table 1 varies. The symbol for volume support structures is a black filled square. If the volume is fragmented for better removability, this can be supplemented by the dimensions shown in Table 2 in the specification field.

If the volume is fragmented for better removability, this can be supplemented by the dimensions shown in Table 2 in the specification field. Column or cone support structures are specified by usage of the corresponding symbol in Table 1. In the following field, the column diameter or the cone diameter at the building platform is listed. In the case of cone support structures, the taper angle εK is also specified. The last field contains dimensions for the arrangement of the individual columns or cones. An example for the specification field is given in Table 2. Tree support structures or complex lattice structures are less relevant to the application of supporting parts, but should still be mentioned here. In principle, they can be described using a specification field, which is why the corresponding symbols are provided (Table 1). However, since many parameters can be varied in the case of tree support structures, the specification field can quickly take on large dimensions.

However, since many parameters can be varied in the case of tree support structures, the specification field can quickly take on large dimensions. Here, not only parameters describing the shape of the individual "trunks" and "branches" and their branching geometry are to be reproduced, but the arrangement pattern of the various individual trees. It is conceivable that each individual support structure tree is different from the others. At most, labels such as "Y-tree" or "IY-tree" could be used for a general characterization, as described by GAN et al [20]. In most cases it will be more appropriate to link the support structures to the engineering drawing as a CAD or STL file.682Table 2: Specifications field for different support types.x,y,z coordinate system of SLM Machine683Lattice support structures can also rarely be described with the help of a few parameters.

Lattice structures of the type "Schwartz Diamond" or "Schoen Gyroid", for example, are relatively easy to describe using the volume fraction in the specification field (Table 2). For more complex lattice structures, however, a reference to a CAD file containing the unit cell is more useful. High complex support structures can also be referenced to a CAD file, this must be indicated in the specification field with the abbreviation "CAD" in the first position of the specification field. In the subsequent field, the CAD file is named by a unique identifier. On a drawing, this can be a file name and the associated path. In a digital environment, a direct link to a database with the reference CAD file is also conceivable.Wall-based support structures: Example drawing entryFigure 7 shows an example component with two different support structure areas. Both areas are connected by a reference line with a rhombus symbol. The rhombus symbol is followed by a specification field.

Both areas are connected by a reference line with a rhombus symbol. The rhombus symbol is followed by a specification field. For the right support structure area, the first line of the specification field indicates that the support structure has a net shape, and a wall thickness of 1.2 mm. The fact that the mesh has four mesh rays and the concentrically running walls have a distance of 3.5 mm can be seen in the third field of the top line. Finally, dimensions for fragmentation are added. The second and third lines of the specification field describe the connection of the support structures with different tooth geometries. Line four of the specification field indicates that the support structure area has a contour wall with a greater wall thickness of 2.2mm and a perforation. The last line describes the component connection of the contour wall. The support structures in the left area are contour support structures with a wall thickness of 0.8 mm. The walls are spaced 2.6 mm apart.

The support structures in the left area are contour support structures with a wall thickness of 0.8 mm. The walls are spaced 2.6 mm apart. The support structures are connected to the component and the building platform with teeth of the same geometry. Moreover there is no difference between hatch and contour walls.Figure 7: Example drawing entry according to Figure 5 and Table 2.Conclusion and outlookIn the present study a method to depict support structures of laser beam melted parts in technical drawings was presented. Therefore, the state of the art was determined, requirements for a useful depiction of supports were identified and a recommondation to depict the support structures in technical drawings was presented. The depiction of support structure areas was set by using cross-hatching in the drawing. To distinguish between different support structure types, a standardised depiction of geometrical information in a specification field was developed.

To distinguish between different support structure types, a standardised depiction of geometrical information in a specification field was developed. The specification field gives a detailed description of the support structure type, the geometry as well as the connection to the part and the building platform. Also uncommon support types like lattice structures or CAD based support structures can be implemented. The forms of depiction developed in this study can serve as a basis for different future support structures. Moreover, additional geometrical information of the support structures can be implemented. An implementation in standardization or other AM processes is also conceivable.684Literature[1][2][3][4][5][6][7][8][9][10][11][12]S. Lammers, T. Lieneke, D.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDeposition Controlled Magnetic Alignment in Iron-PLA Composites Nathan D. Watson, and Paris Von LocketteDepartment of Mechanical and Nuclear Engineering, The Pennsylvania State University, University Park, PA, 16802Abstract By manipulating the print plane, infill direction, and geometry of Fused Filament Fabricated (FFF) iron-PLA composite parts, the alignment of their magnetic axes can be influenced. FFF printing allows control of deposition direction, which affects the arrangement of the iron within the composite part in ways that induce preferred magnetic orientation, the so- called easy axis. Qualitative results show the direction of deposition of the composite iron-PLA filament has significant effects on the response of the printed parts to an external magnetic field.

Qualitative results show the direction of deposition of the composite iron-PLA filament has significant effects on the response of the printed parts to an external magnetic field. Results further show that across different geometries, the easy axis of a printed part can be prescribed by setting the print plane and infill direction parallel to the desired orientation. Expected part geometry effects, along with the print plane and infill influences, suggest the phenomenon can be modeled using multi-scale demagnetizing field theories to print magneto-sensitive devices that can perform localized, controlled actuation in a uniform magnetic field.IntroductionAdditive manufacturing (AM) is a process used to create 3D parts by depositing materiallayer by layer. Fused Filament Fabrication (FFF) uses a heated extruder to melt and deposit material in a 2D pattern to create each layer.

Fused Filament Fabrication (FFF) uses a heated extruder to melt and deposit material in a 2D pattern to create each layer. AM offers advantages over other manufacturing techniques by providing free complexity and a unique control over how the material is structured in a 3D part. This allows the fabrication of relatively inexpensive, highly customized low volume production. This allows for specific customization for individual applications or people.Recent advances have led to the development of 3D printed structures with embedded electromagnetically sensitive inclusions for a range of applications including communications, locomotion, actuation, and control1,2,3,4. In these works, the inclusions were biased toward a preferential direction using an external, controlled, electromagnetic field generator that physically reorients the inclusions either before, during, or after deposition of the initial feed stock.

The processing therefore necessarily requires additional equipment, complexity and ultimately expense beyond simple deposition of the existing stock material.Previous work has shown that direction of deposition can affect a printed part’smechanical response to loading5. By controlling the direction of deposition, the extruded filament which make up a printed part can have a major axis on which the fibers lay. In magnetism the concept of a major axis plays an important role in how a part responds in a magnetic field. Magnetic materials may develop a so-called easy axis along which they are more susceptible to external fields, e.g. along the length of a nail or from the south to the north poles in a bar magnet. This work seeks to examine this phenomenon.C. Huber, C. Abert, F. Brucknet et al.

along the length of a nail or from the south to the north poles in a bar magnet. This work seeks to examine this phenomenon.C. Huber, C. Abert, F. Brucknet et al. imbedded NdFeB particles in extrudable plasticfilament and were able to use a dual feed extruder to mix NdFeB composite with a base plastic to1145vary the weight percentage throughout a printed part6. By changing the weight fraction at different parts of the geometry, they were able to design the remanence field around the part. While they mainly realized the effects of the weight fraction in the part they didn’t research how other printing parameters effected the magnetic properties of the parts.L. Bollig, M.Patton, G.Mowry et al. looked at how certain print parameters changed the magnetic properties of printed parts using Proto-Pasta Iron PLA composite7. They measured the response of test samples in the VSM after changing the infill percentage, angle of infill pattern, and number of parameter walls.

They measured the response of test samples in the VSM after changing the infill percentage, angle of infill pattern, and number of parameter walls. While they were able to find differences based on how the part was oriented in the VSM, they didn’t find how infill can play a major role in the magnetic properties of a part.By manipulating some of the printing parameters of a printed part, the structure of thedeposited fibers can be arranged in a way to make an easy axis for the part which causes the part to physically align with a uniform magnetic field. Both infill direction and print orientation create strong effects on the easy axis and by manipulating each, two constraining moments can be applied on a part to align a part in a particular orientation in a magnetic field. By setting a particular print orientation, an easy plane is created. The part orients itself in a way that makes the print plane parallel to the field.

By setting a particular print orientation, an easy plane is created. The part orients itself in a way that makes the print plane parallel to the field. This allows the part to rotate freely on that plane with no opposing magnetic moment. By adding an infill direction perpendicular to the print plane, a second easy plane is created, and where those planes intersect is the easy axis, or the orientation the part takes in a uniform magnetic field.Just as fibers want to align in a magnetic field in a part, the same principal can be used,but instead of moving the fibers in the part to align with the field, the 3D printer easily allows for the arrangement of fibers and lets the user define the orientation of the fibers before the print.

As magnetic fibers are placed in a magnetic field; a magnetic moment is created in order to reduce the demagnetizing field and in turn reduce the Zeeman energy of the system.MethodsWhen placed into a uniform magnetic field, the parts printed would align to the predefined easy axis. This meant that depending on the print parameters and print orientation set for printing a particular item like a cylinder, defined how the cylinder would align in the magnetic field. Cylinder A from figure 1 was printed on its side making the print plane along the axis of the cylinder. This allows the cylinder align itself parallel with the field. Cylinder B from figure 1 was printed vertically on the print bed making the print plane perpendicular to the cylinder.

This allows the cylinder align itself parallel with the field. Cylinder B from figure 1 was printed vertically on the print bed making the print plane perpendicular to the cylinder. This makes the cylinder orient itself horizontally in the magnetic field.ABFigure 1: Cylinder [A] printed horizontally cylinder [B] printed vertically in a uniform magnetic field, being held in position by the magnetic moment on the print.1146In order to test the magnetic response to the effects of infill direction and print planeorientation, samples were made to isolate each. By printing flat cylinders and rotating the cylinders along its axis, the geometry of the part was taken out of the test because the sample maintains its geometry at each angle the sample is tested at. To isolate the infill of the sample, the cylinder was printed flat on the print bed so that the print plane would be at the same orientation relative to the field as the cylinder is rotated.

To isolate the infill of the sample, the cylinder was printed flat on the print bed so that the print plane would be at the same orientation relative to the field as the cylinder is rotated. The infill direction in the sample was printed so that each deposited fiber was parallel and as seen in figure 3A. The sample designed to isolate the print plane was printed on its edge with the infill running from the base of the cylinder to the top. This way as the cylinder rotates about its axis the print plane changes while the infill stays perpendicular to the field as seen in figure 3B. A control was also printed, but instead the infill is comprised of concentric circles which makes the sample axially symmetric. The samples were tested in a vibrating sample magnetometer (VSM) as seen in figure 2.

The samples were tested in a vibrating sample magnetometer (VSM) as seen in figure 2. The samples were 8mm in diameter and 2mm tall.Figure 2: VSM machine testing a sample in the transverse positionThe samples were printed on an Ultimaker 3 using Proto-Pasta iron PLA which is 45%iron by weight. [Proto-Pasta] The samples were then testing on a two tesla VSM in the transverse direction, rotating the parts on the axis of the cylinder. The samples were tested at 5 different angles, 0 degrees being where the field lines produced by the VSM is parallel to either the print plane or infill direction and 90 degrees being perpendicular. At each angle the samples were tested from -18900 Oe to 18900 Oe. The VSM is able to find at what field the sample saturates at each angle.ResultsWhen tested in the VSM machine, the samples showed that 3D printed parts can exhibitanisotropic properties by manipulating the printing parameters.

For both the parallel infill sample and the print plane sample, the linear portion of the curve approached saturation faster in the orientation with the infill or print plane parallel to the field as seen in figure 3. Figure 4 shows the slope of the curves of the samples at each angle at 1000 Oe, which is at a linear portion of the curve. The sample with concentric circle infill did not show significant changes in the magnetization based on the angle of the sample in the VSM machine.The saturation of the samples at each angle were normalized to 1 in order to account for the slight variations from the slight change in position in the VSM as the angle changes as well as the differences in mass of the samples. As the sample in the VSM machine rotated, the position between the magnetic pick up coils would change causing a difference in measured saturation while the saturation is not dependent on the position or angle of the magnetic material in the VSM machine7.

The inconsistencies of the printing process also caused the samples to have different masses despite having an identical bulk shape. The more magnetic material in the VSM machine would raise the saturation of the sample. By normalizing the saturation to one, it also corrects for the inconsistencies of the mass and the possible slight inconsistencies in the weight fraction of the Proto-Pasta iron PLA.1147ABCABCFigure 3: Hysteresis Curve of the [A] print plane, [B] infill direction, and concentric infill samples at 0, 30, 60, and 90 degrees.0.000240.000230.00022µ0.000210.00020.000190.00018Print PlaneParallel InfillCircular Infill0 Degrees30 Degrees60 Degrees90 DegreesFigure 4: Slope of the normalized hysteresis curve at each angle at 1000 Oe.1148Analyses and DiscussionThe authors believe two factors, well known and important to magnetization response,produce the results shown in this work.

First is the concept of the easy axis, a geometry- dependent direction toward which induced magnetization is biased. In this work the authors contend that the infill and print plane are able to create easy axes apart from the bulk part geometry. The orientation of the easy axis is critical to actuation, e.g. magnetic torques acting on the bulk part. When subjected to an applied magnetic field, a magnetic body will tend to align its easy axis with that external field direction. The underlying mechanism of alignment is minimization of the Zeeman energy9, given as𝑯𝑯𝑒𝑒𝑒𝑒𝑒𝑒,(1)𝑈𝑈 = −𝜇𝜇0𝒎𝒎 ∙ 𝑯𝑯𝑒𝑒𝑒𝑒𝑒𝑒 where object is exposed. While the easy axis is not necessarily collinear with the easy axes direction.

While the easy axis is not necessarily collinear with the easy axes direction. A thorough discussion of this phenomena is beyond the scope of this work but may be found in any text on magnetic behavior of materials9.is the magnetization of the object in question andis the external field to which the,is biased toward𝑯𝑯𝑒𝑒𝑒𝑒𝑒𝑒𝒎𝒎𝒎𝒎𝒎𝒎From (1) it is clear that the object will tend to have its internal magnetization, and henceits easy axis, aligned with the external field. This tendency produces the well-known magnetic torque exerted on magnetic objects in an external field that has been used as an actuation mechanism in magneto-active composites10,11,12,13.

This tendency produces the well-known magnetic torque exerted on magnetic objects in an external field that has been used as an actuation mechanism in magneto-active composites10,11,12,13. As seen in figure 1, the bulk actuation of the printed part follows the easy axes created by preferential infill or print plane orientation, which may be in opposition to the easy axes expected from the bulk geometry.The second important concept is that of the stray field, the external magnetic field generated by a magnetically susceptible body when the body is exposed to an external field. This stray field is important because it affects neighboring magnetic bodies by changing the external field local to them.

This stray field is important because it affects neighboring magnetic bodies by changing the external field local to them. In the case of soft magnetic material, such as the iron-PLA composites used herein, the internal magnetization can be expressed by(2)𝒎𝒎 = 𝜇𝜇𝑯𝑯𝒆𝒆𝒆𝒆𝒆𝒆 where magnetic bodies surrounding the body in question such that the magnetization of the body in question can be expressed asis the magnetic permeability of the composite. Stray fields,, arise from all discrete𝑯𝑯𝑠𝑠𝜇𝜇(3)𝒎𝒎 = 𝜇𝜇�𝑯𝑯0 + ∑ 𝑯𝑯𝑠𝑠𝑖𝑖 sums the contribution of the where stray fields of all discrete, magnetically susceptible bodies in the system to the external field at 𝑖𝑖 the body in question.

From (3) it is clear that if we ignore stray fields, the magnetization of the body, in the small field region, would nominally scale linearly with inclusion of stray fields complicates matters, allowing the size, shape, orientation, and composition of neighboring magnetic bodies to influence magnetization response. In this work, the stray fields differentially affect the magnetization response of our printed geometries by𝑖𝑖�is the external field applied to the body in question and𝑯𝑯0. However, the‖𝑯𝑯0‖1149diminishing the local field in cases where parallel planes are normal to the measurement axis and enhancing cases where parallel planes are parallel to the measurement axes, yielding the results of figure 5.Figure 5: Iron PLA fibers in a magnetic field.

position of Ti/TiC Composite Coatings on Implant Structures using Laser Engineered Net ShapingG.D. Janaki Ram, Y. Yang, and B.E. Stucker Department of Mechanical and Aerospace Engineering, Utah State University Logan, UT 84322-4130, USA Reviewed, accepted September 5, 2007AbstractA new method of depositing hard and wear resistant composite coatings on metal-on- metal bearing surfaces of titanium implant structures is proposed and demonstrated. The method consists of depositing a Ti/TiC composite coating (~ 2.5 mm thick) on titanium implant bearing surfaces using Laser Engineered Net Shaping (LENS®). Defect-free composite coatings were successfully produced at various amounts of the reinforcing TiC phase with excellent interfacial characteristics using a mixture of commercially pure Ti and TiC powders.

The coatings consisted of a mixture of coarser unmelted/partially melted (UMC) TiC particles and finer, discreet resolidified (RSC) TiC particles uniformly distributed in the titanium matrix. The amounts of UMC and RSC were found to increase with increasing TiC content of the original powder mixture. The coatings exhibited a high level of hardness, which increased with increasing TiC content of the original powder mixture. Fractographic studies indicated that the coatings, even at 60 vol.% TiC, do not fail in a brittle manner. Various aspects of LENS® deposition of Ti/TiC composite coatings are addressed and a preliminary understanding of structure-property-fracture correlations is presented.

IntroductionMaterials used for biomedical implant devices must satisfy a variety of property demands, which are often mutually exclusive. Further, different parts of a device demand different material properties. These factors often make it difficult to manufacture a medical device using a single material. Many implant devices, such as knee replacements, involve a relative motion between constituent parts while in service. In such situations, polyethylene-on- metal bearing surfaces are currently being widely used due to the forgiving nature of the combination to the individual biomechanical nuances of joint recipients. Studies, however, show that, although polyethylene is biologically inert in the body as a whole, microscopic particles of polyethylene resulting from wear can possibly be toxic if encountered in large amounts and lead to osteolysis [1-3].

This has led to a growing interest in metal-on-metal bearing surfaces for implants, especially in applications involving a large contact area [4-5].The metal-on-metal approach, however, requires the bearing surfaces to be extremely hard and wear resistant. Fabrication of an entire implant device from materials hard enough to satisfy these requirements is not plausible, as such hard materials would be too brittle to use. A more plausible approach would be to use a suitable hard surface coating on the bearing surfaces of the implant. This approach allows one to manufacture an implant device from well-established the surface property biocompatible materials with desired bulk characteristics, while requirements are met by using a biocompatible hard surface coating.While the surface coating approach is certainly advantageous, it presents a number of challenges related to coating materials and coating methods.

Some of the issues related to coating materials are: i) the coating material should be biocompatible, ii) the coating should not be brittle or prone to spalling, iii) the coating should be adequately hard and wear resistant, and iv) the coating material should be metallurgically compatible with the substrate material. While most of the candidate coating materials, such as carbides, borides and nitrides, easily meet the hardness and wear resistance requirements, all of them are inherently brittle and are often not metallurgically compatible with metallic implant substrate materials. Further, it is difficult to apply robust, residual stress free, single-phase non-metallic coatings due to their inherent brittleness and propensity to spalling when deposited in thin layers. For example, application of single-phase TiC coating on a Ti alloy substrate using a laser-based metal deposition process was shown to result in gross cracking [6].

For example, application of single-phase TiC coating on a Ti alloy substrate using a laser-based metal deposition process was shown to result in gross cracking [6]. These considerations led us to consider using a composite coating constituting a metallic matrix material and a hard non-metallic material, in the form of uniformly dispersed particles in the matrix. In this way, one can produce a coating with all the desired characteristics ─ the softer metallic matrix material provides toughness and metallurgical compatibility/bonding with the substrate, whereas the hard non-metallic reinforcing material ensures adequate overall coating hardness and wear resistance.Another important issue is the method of coating application.

Some of the challenges in this regard are: i) the coating method should be capable of producing sound, fully-dense composite coatings, ii) the coating method should be capable of achieving metallurgical bonding between the coating and the substrate material in order to ensure adequate bond strength, iii) the coating method should use minimal heat input in order not to cause excessive microstructural damage to the substrate material, iv) the coating method should allow precise control over the coating application area to facilitate coating application just where needed on the implant device, and v) the coating method should have a material deposition rate capable of achieving adequate coating thickness to resist spalling within a reasonable amount of time.

Further, it may be desirable, depending on the specific application requirements, to produce a coating with a spatial2variation in the amount or type of second phase particles to facilitate application of functionally graded coatings for achieving superior coating performance. Gradual increase in second phase amount provides a smooth microstructural and/or chemistry transition, allowing coatings to be deposited without cracking problems. Most conventional coating processes, such as thermal spray processes, diffusion coating processes, arc welding-based cladding processes, vapor phase processes, and others, are not quite suitable in the current context, as each of them fail to satisfy one or more of the above requirements.Laser-based metal deposition processes satisfy all of the above requirements and are thus quite promising for producing composite coatings on implant surfaces.

In the current work, one of the commercially available laser-based metal deposition processes, Laser Engineered Net Shaping® (LENS), has been used. LENS® is a novel additive manufacturing process developed by Sandia National Laboratories for fabrication of fully-dense three-dimensional (3D) metallic objects. A schematic of the LENS® process is shown in Fig.1. Initially, a three-dimensional CAD model of the component to be built is generated and a computer program slices the model into a number of horizontal cross-sections or layers. These cross-sections are systematically created from bottom to top producing a three-dimensional object fully automatically. The deposition process begins with directing a focused Nd-YAG laser beam onto a substrate placed on the build platform capable of computer-controlled motion. The laser generates a small molten pool (typically 0.25-1 mm in diameter and 0.1-0.5 mm in depth) on the substrate.

The laser generates a small molten pool (typically 0.25-1 mm in diameter and 0.1-0.5 mm in depth) on the substrate. Powdered material at a selected rate is injected directly into the melt pool using a powder feeder. The molten pool solidifies rapidly as the laser beam moves away, forming a thin track of solidified metal welded to the material below along the line of laser scanning. A layer is generated by a number of consecutive overlapping tracks. After each layer is formed, the laser head, along with the powder delivery nozzle, moves upward by one layer thickness and the subsequent layer is generated. This process is repeated until the part is complete. The deposition process occurs inside an enclosed chamber filled with argon to prevent oxidation of the liquid metal. Although the process was originally developed for fabricating free-standing 3D parts, the process lends itself to layer-by-layer deposition of coatings of desired thickness.Fig.1.

Although the process was originally developed for fabricating free-standing 3D parts, the process lends itself to layer-by-layer deposition of coatings of desired thickness.Fig.1. Schematic of the LENS® process.3The LENS® process has successfully been applied to fabrication of engineering components in materials like stainless steels, tool steels, titanium alloys, and superalloys [7-11]. Attempts have also been made to produce composite materials and functionally graded materials. For example, Banerjee et al. [12] produced a gradual transition in LENS® deposit composition and microstructure from binary Ti–8Al to Ti–8Al–20V within a length of 25 mm using elemental powders. Liu and DuPont [6] produced functionally graded Ti/TiC composite deposits using the LENS® process, employing separate powder feeders for introducing Ti and TiC powers in desired quantities.

Liu and DuPont [6] produced functionally graded Ti/TiC composite deposits using the LENS® process, employing separate powder feeders for introducing Ti and TiC powers in desired quantities. The investigators also fabricated TiC/TiAl composites, using preheated titanium substrates to prevent solid-state cracking in the deposit due to thermal stresses [13]. Similarly, Banerjee et al. [14] produced Ti-6Al-4V/TiB composite parts employing a blend of pure pre-alloyed Ti-6Al-4V and elemental B powders. In contrast, very few attempts have been made to utilize the process for the purpose of depositing a coating. Kumar and Stucker [15] have successfully deposited CoCrMo coating on a porous Ta substrate using the LENS® process employing Zr as an interlayer. While work by Liu and DuPont [6,13], and Banerjee et al.

While work by Liu and DuPont [6,13], and Banerjee et al. [14] are closely relevant in the current context, no reports are available on the use of the LENS® process for deposition of wear resistant Ti/TiC composite coatings on implant structures.The current work explores the possibility of depositing Ti/TiC composite coatings on titanium implant surfaces using the LENS® process. Studies were directed to examine the effect of TiC volume fraction on coating formation, microstructure, and hardness, and to develop a preliminary understanding of structure-property-fracture correlations.2. Experimental WorkTi/TiC composite coatings 15x15x2.5 mm were produced using an Optomec 750 LENS® machine equipped with a 500 W Nd-YAG laser. The oxygen level in the deposition chamber was kept below 5 ppm during processing to avoid oxygen and nitrogen contamination in the deposit.

The oxygen level in the deposition chamber was kept below 5 ppm during processing to avoid oxygen and nitrogen contamination in the deposit. The coating material consisted of a mixture of commercially pure (CP) Ti and TiC powders (– 100/+325 mesh size (45-150 µm)). Prior to deposition experiments, Ti and TiC powders in selected quantities on a volume basis were thoroughly mixed in a blender and the mixture was fed into a powder feeder.Table 1. LENS® parameters used for depositing Ti/TiC coatings.Parameter Level 165 Laser power (W) Laser traverse speed (mm/s) 17 Powder feed rate (g/minute) 4 Layer thickness (mm) Hatch spacing (mm) Scan orientation (°)0.12 0.25 0, 90, 0Initially, powder mixtures were prepared containing 20% and 40% TiC by volume. Coatings produced using these powder mixtures will be referred to as Ti/20TiC and Ti/40TiC. For these experiments, a (CP) Ti wrought-processed sheet of 1 mm thickness was used as the substrate (substrate dimensions: 30 x 30 mm).

For these experiments, a (CP) Ti wrought-processed sheet of 1 mm thickness was used as the substrate (substrate dimensions: 30 x 30 mm). The parameters employed for coating deposition are listed in Table 1. These parameters were found to result in sound coatings and were arrived at after preliminary deposition experiments involving the use of laser power, traverse speed,4powder feed rate, layer thickness, and hatch spacing at various levels. Scan orientation was changed (0°, 90°, 0°) for every layer in order to minimize residual stress build up.After examining the microstructures of the Ti-20TiC and Ti/40TiC coatings, further coating experiments were performed with increased TiC content (60 vol.%) in the powder mixture. The coating produced with this powder mixture will be referred to as Ti/60TiC. For these experiments, a CP Ti wrought-processed plate 3 mm think was utilized as the substrate (substrate dimensions: 100 x 50 mm).

For these experiments, a CP Ti wrought-processed plate 3 mm think was utilized as the substrate (substrate dimensions: 100 x 50 mm). The process parameters utilized for these experiments are the same as those reported in Table 1.Sections cut from LENS® deposits were prepared for microstructural examination following standard metallographic practices. Polished specimens were etched with Kroll’s reagent (95 ml H2O, 3 ml HNO3, and 2 ml HF). Microstructural studies were performed using an Axiovert 100A inverted light microscope and a Hitachi S 4000 Scanning Electron Microscope (SEM) equipped with a field emission gun. In order to assess the fracture behaviour of coatings, coatings were fractured under impact loading and the fracture surfaces were examined under SEM. Hardness tests were performed on transverse sections of the coatings using a Vickers microhardness tester at a load of 300g. At least five tests were performed on each specimen and the average values were reported.3.

At least five tests were performed on each specimen and the average values were reported.3. Results and Discussion 3.1 Coating MicrostructuresThe SEM microstructures of the Ti/TiC coatings (transverse section) produced using 20, 40, and 60 vol.% TiC powder mixtures are shown in Fig.2. All the coatings were found to be free from gross cracking, porosity and other defects. Individual layers/tracks in the coatings, characteristic of LENS® deposited materials, are apparent in these microstructures. The coating/substrate interface in all cases was observed to be free from defects, as can be seen in Fig.2d. Observations thus show that the LENS® process can produce sound Ti/TiC coatings over a wide range of TiC contents. In fact, successful LENS® deposition of a functionally graded Ti/TiC coating on a Ti-6Al-4V substrate, where the TiC content was increased from 0 to 95 vol.% (in the powder mixture), was reported by Liu and DuPont [6].

In the present study, coatings with higher than 60 vol.% TiC were not attempted for two reasons: i) use of large amounts of TiC can result in the formation of continuous carbide networks (as observed by Liu and DuPont [6]), which render the coatings brittle, and ii) deposition of coatings with a very high TiC content directly onto a Ti substrate can result in cracking at the substrate/coating interface (the situation is different when depositing a functionally graded coating, where the TiC content is gradually increased). In this context, Liu and DuPont [6] reported cracking problems in Ti/TiC composite coatings with 80 vol.% TiC, when the coating was directly deposited onto a Ti-6Al- 4V substrate.The microstructures shown in Fig.2 reveal the presence of quite a few irregularly-shaped, uniformly-distributed unmelted TiC particles (hereafter referred to as UMC), whose size ranged from 30 to 90 µm, in all the coatings.

The amount of UMC can be seen to increase with increasing TiC content in the powder mixture, while the particle size appeared to be the same in all cases. Liu and DuPont [6] also reported similar observations. Quantitative image analysis on coating microstructures showed that the volume fraction of UMC in the three coatings was approximately 50% of the TiC content used in respective powder mixtures (Ti/20TiC, Ti/40TiC, and Ti/60TiC coatings contained 10.2, 19.5, and 28.6 vol.% of UMC, respectively). These5particles result from incomplete melting of the TiC powder particles originally introduced into the melt pool, whose size ranged from 45 to 150 µm.(a)(b)(c) (d) Fig.2.

These5particles result from incomplete melting of the TiC powder particles originally introduced into the melt pool, whose size ranged from 45 to 150 µm.(a)(b)(c) (d) Fig.2. SEM microstructures of Ti/TiC composite coatings: (a) Ti/20TiC, (b) Ti/40TiC, (c) Ti/60TiC, (d) Ti/60TiC (coating/substrate interface shown by arrow).A comparison of the amounts and size of TiC originally introduced in the powder mixtures with the volume fractions and size of UMC in the coating microstructures shows that TiC particles undergo considerable melting/dissolution during LENS® processing. It may be noted that the melting point of TiC (3160°C) is significantly higher than that of Ti (1668°C). In6spite of this, partial melting of TiC can occur during LENS® processing because of the higher laser absorption coefficient and lower thermal conductivity of TiC in relation to Ti [6]. Further, TiC particles can also experience some amount of dissolution in liquid titanium.

Further, TiC particles can also experience some amount of dissolution in liquid titanium. The extent of TiC melting/dissolution is basically governed by the magnitude of incident laser energy and laser/TiC as well as liquid metal/TiC interaction times. Since all the three coatings were deposited using the same process parameters, it is only natural that the volume fraction of UMC increases with increased TiC content in the initial powder mixture.(a) (b)(c) Fig.3. High magnification SEM microstructures of Ti/TiC composite coatings: (a) Ti/20TiC, (b) Ti/40TiC, and (c) Ti/60TiC. The plate-like features seen in (a) and (b) are coarse α-Ti plates.Examination of coating microstructures at higher magnification revealed numerous finer TiC particles (about 1-4 µm) in all the coatings (Fig.3). These particles were confirmed to be non-stoichiometric TiC (TiC0.65) by Liu and DuPont [6] based on X-ray diffraction studies.

These particles were confirmed to be non-stoichiometric TiC (TiC0.65) by Liu and DuPont [6] based on X-ray diffraction studies. These finer carbide particles were found to be generally discrete and uniformly distributed in the7coating microstructure. A closer examination reveals that the non-stoichiometric TiC phase occurs in two different morphologies: i) slightly larger, dendritic-type TiC particles, and ii) very fine equiaxed TiC particles. The size and amount of dendritic type carbide particles appeared to increase with increase in TiC content in the powder mixture (Fig.3). The matrix in all cases consisted of martensitic α-Ti in a widmanstatten pattern. However, α-Ti plates appeared to be significantly coarser in Ti-20TiC and Ti-40TiC coatings compared to Ti-60TiC (Fig.3).Development of microstructure in Ti/TiC composite coatings can be understood in the following manner.

As noted earlier, the TiC particles present in the powder mixture undergo considerable melting/dissolution during LENS® processing. This results in carbon enrichment, consistent with the extent of dissolution, in the liquid metal. Solidification of this carbon-rich Ti melt takes the following sequence: L → L + primary TiC → β-Ti+TiC eutectic [6,13]. Solidification starts with formation of primary TiC particles. According to the Ti-C phase diagram, C concentrations from approximately 0.5 to 16 wt.% C will form primary TiC during solidification [16]. Primary TiC appears as dendritic particles in the microstructure. As this happens, the carbon level in the remaining liquid comes down and eventually primary TiC formation would stop. When the eutectic temperature is reached (~ 1648°C), the L → β-Ti+TiC reaction takes place terminating the solidification process. The eutectic TiC would be finer and equiaxed compared to primary dendritic TiC.

The eutectic TiC would be finer and equiaxed compared to primary dendritic TiC. Thus at the end of solidification, the microstructure contains UMC particles, primary dendritic TiC particles, eutectic TiC, and β-Ti. Thus both primary dendritic TiC and eutectic equiaxed TiC form during solidification from a C rich Ti melt. These TiC particles will be referred to as resolidified carbide (RSC) particles. During cooling after solidification very little change in carbide microstructure occurs [6]. However, when the β-transus temperature (~900°C) is reached, β-Ti transforms to α-Ti. This transformation can occur either diffusionally or martensitically or a combination of the two, depending on the cooling rates [17]. Because of the rapid cooling rates in the LENS® process (in the range of 104 to 105 °C/s during solidification [18], but likely lower during the solid-state transformation discussed here) this transformation is expected to occur almost entirely by martensitic transformation.

Thus, the coating microstructure at room temperature consists of UMC and RSC in a matrix of widmanstatten martensitic α-Ti.The observed differences in the microstructures of the three coatings can be explained as follows. The size and amount of primary dendritic TiC particles was observed to increase with increasing TiC content in the powder mixture (Fig.3). This should be a result of increased C content in the Ti melt. The estimated UMC volume fractions in the coatings indicate that approximately 50% of the TiC introduced in the powder mixtures melts/dissolves during LENS® processing in all cases. Consequently, C content in the liquid metal would increase with increasing TiC content in the initial powder mixture. An increase in C content in the liquid metal would mean more C availability for primary TiC formation. This results in more and larger primary TiC particles in the microstructure, as observed in the Ti/60TiC coating (Fig.3c).

This results in more and larger primary TiC particles in the microstructure, as observed in the Ti/60TiC coating (Fig.3c). Further, higher carbon contents in liquid metal result in higher constitutional super cooling levels, which favor formation of primary Ti particles with a strong, pronounced dendritic shape [6,13]. This explains why primary TiC particles are more dendritic in appearance in Ti/60TiC coating than those in the other two coatings. It should be noted that the amount of eutectic TiC remains more or less the same in all cases, as primary TiC formation continues as long as the C content in the liquid is more than 0.5 wt. %, leaving the same amount of C in the remaining liquid before the eutectic reaction takes place.8The observed variations in matrix microstructure among the three coatings can be explained considering the differences in substrate dimensions.

Ti/20TiC and Ti/40TiC coatings were deposited on a 1 mm think Ti sheet (30 x 30 mm coupons), while the Ti/60TiC coating was deposited on a 3 mm thick Ti plate (100 x 50 mm coupons). Because of the larger substrate thickness and size (i.e., greater volume of heat sink), cooling rates would be higher in Ti/60TiC coating compared to the other two coatings. Substrate dimensions also determine the amount of heat built up during coating deposition. Because of the smaller substrate dimensions, there can be more heat built up in the Ti/20TiC and Ti/40TiC coatings than in the Ti/60TiC coating, although all the coatings were deposited using the same heat input. Consequently, these coatings experience significantly lower cooling rates, which affect β-Ti to α-Ti transformation.

Consequently, these coatings experience significantly lower cooling rates, which affect β-Ti to α-Ti transformation. While β- Ti to α-Ti transformation can still be predominantly martensitic, the reduced cooling rates in these cases can result in some amount of diffusional transformation, leading to the formation of coarser, parallel plates of diffusional α.Another factor that influences the matrix microstructure is multiple thermal cycling. LENS deposited materials experience a complex thermal history in a manner very similar to multi-pass weld deposits. The heat energy introduced during depositing a layer can reheat previously deposited material into various phase fields. The thermal cycle experienced can be different at each point in the reheated material depending on its distance from the heat source. The multiple thermal cycles thus can result in coarsening of α-Ti plates in the matrix microstructure.

The multiple thermal cycles thus can result in coarsening of α-Ti plates in the matrix microstructure. These effects are expected to be more pronounced in Ti/20TiC and Ti/40TiC coatings due to a larger amount of heat built up expected in these cases.3.2 Coating HardnessThe results of microhardness tests on coatings and a CP Ti substrate are given in Table 2. Hardness measurements in the coatings were performed in matrix regions (i.e., away from UMC particles in the coating microstructures). Therefore, the reported coating hardness values represent the hardness of the Ti matrix with finer RSC particles. The bulk hardness levels of the coatings are expected to be higher than the reported values, as UMC particles also contribute to coating hardness during bulk hardness testing.Table 2.

The bulk hardness levels of the coatings are expected to be higher than the reported values, as UMC particles also contribute to coating hardness during bulk hardness testing.Table 2. Results of Vickers microhardness tests (average of 5 measurements).Coating Ti/20TiC Ti/40TiC Ti/60TiC Ti substrateHardness (VHN) 410 760 1190 195As can be seen in Table 2, all the coatings exhibited significantly higher hardness values when compared to the CP Ti substrate due to the presence of carbide particles. Coating hardness can be seen to increase with increasing TiC content in the original powder mixture. This correlates well with microstructural observations. As noted previously, the amount and size of RSC particles increased from Ti/20TiC to Ti/40TiC to Ti/60TiC. Consequently, Ti/40TiC and Ti/60TiC coatings exhibited relatively higher hardness levels when compared to Ti/20TiC coating.

Consequently, Ti/40TiC and Ti/60TiC coatings exhibited relatively higher hardness levels when compared to Ti/20TiC coating. The results thus show that the hardness of titanium implant surfaces can be substantially increased by LENS® depositing a Ti/TiC composite coating. It may be noted that even higher9hardness levels are achievable in the coatings by simply increasing the TiC content in the coatings. However, there may be attendant problems as discussed in Section 3.1.3.3 Fracture Behaviour of Ti/TiC Composite CoatingsTi/TiC composite coatings were fractured by applying an impact load to assess the fracture behaviour of the coatings. The SEM fractographs of a Ti/60TiC coating are shown in Fig. 4. All the coatings showed fractured UMC particles on the fracture surfaces (Fig.4a). It may be noted that there were no completely removed UMC particles on the fracture surface, which indicates that the UMC/matrix interfaces are strong (Fig.4a and Fig.4b).

It may be noted that there were no completely removed UMC particles on the fracture surface, which indicates that the UMC/matrix interfaces are strong (Fig.4a and Fig.4b). Examination of fracture surfaces, however, shows that UMC particles are fractured in a brittle manner during the fracture process (Fig.4c). Brittle fracture of UMC particles is no surprise considering the inherent brittleness of TiC. Importantly, the matrix regions (regions surrounding UMC particles, in Fig.4b and Fig.4c) exhibited ductile fracture features with well-defined dimples, as can be seen in Fig.4d. Some particles, presumably RSC particles, were seen inside the dimples, which suggest these particles served as void nucleation sites (Fig.4d). Nucleation of voids at hard second phase particles is quite common.

Nucleation of voids at hard second phase particles is quite common. Studies thus confirm that LENS® deposited Ti/TiC composite coatings do not fail in a brittle manner, even at relatively high levels of TiC content.The presence of coarse UMC particles in the coating microstructure appears undesirable from many angles. First, UMC particles fracture in a brittle manner, leading to an overall reduction in coating toughness. Coatings can be made more ductile by eliminating the UMC particles. Second, UMC particles offer resistance to matrix shrinkage during solidification, which may result in generation of undesirable tensile residual stresses at the UMC/matrix interfaces. Third, UMC particles present favorable sites for formation of defects in the coating. Fourth, UMC particles may be removed early in the wear process, which may lead to significant third body wear.

Third, UMC particles present favorable sites for formation of defects in the coating. Fourth, UMC particles may be removed early in the wear process, which may lead to significant third body wear. Finally, even without any contribution from UMC particles, reasonably high hardness levels are achievable in Ti/TiC coatings. Further studies involving metal-on-metal wear tests on Ti/TiC coatings are necessary to assess their wear performance and to see if the UMC particles are really undesirable.As demonstrated by Liu and DuPont [6], use of higher heat inputs facilitate more complete melting/dissolution of TiC particles in the powder mixture and, therefore, help eliminate or minimize UMC particles in the coating microstructure. However, use of higher heat inputs alone may not completely eliminate UMC particles, especially in coatings with a high level of TiC.

However, use of higher heat inputs alone may not completely eliminate UMC particles, especially in coatings with a high level of TiC. Also, use of higher heat inputs can have other undesirable effects, such as formation of coarse microstructures and/or generation of excessive residual stresses, which may lead to cracking/delamination of the coating and affect coating properties. Another factor that assumes importance in this context is TiC particle size. Use of relatively finer TiC particles in the powder mixture can help eliminate or minimize UMC particles. Finer TiC particle sizes facilitate greater absorption of incident laser energy leading to greater melting of the particles. Finer particles also dissolve to a greater extent in the molten pool. Therefore, it should be possible to produce a Ti/TiC composite coating without a large number of coarse UMC particles by appropriately controlling the input TiC particle size without unduly increasing the heat input.

However, care must be exercised to avoid use of very fine powders, as very fine powder sizes present difficulties with uniform powder feeding in the LENS® system.10(a) (b)(c) (d) Fig.4. SEM fractographs of Ti/60TiC coating: (a) Fracture surface at a low magnification showing fractured UMC particles, (b) Fracture surface at a higher magnification showing brittle fractured UMC particles (note ductile fracture features in the matrix region), (c) Brittle fractured UMC particle at a higher magnification, and (d) Dimpled rupture features at a higher magnification in the matrix region.4. ConclusionsMetallurgically sound Ti/TiC composites can be deposited on titanium implant surfaces using Laser Engineered Net Shaping process. Coatings can be deposited over a wide range of TiC contents.Ti/TiC coatings show uniformly distributed partially melted TiC particles (UMC) and fine resolidified (RSC) TiC particles.

Coatings can be deposited over a wide range of TiC contents.Ti/TiC coatings show uniformly distributed partially melted TiC particles (UMC) and fine resolidified (RSC) TiC particles. The amounts of UMC and RSC increase with increasing TiC content in the powder mixture.Coating hardness increases with increasing TiC content in the powder mixture due to increased carbide precipitation.11Fracture in Ti/TiC composite coatings occurs in a ductile manner. UMC particles fracture in a brittle manner, making their presence undesirable in the coating microstructures. UMC particles can be eliminated or minimized by controlling the input TiC particle size and by using higher heat inputs.AcknowledgementsThe authors would like to thank Daniel Justin and MedicineLodge, Inc. for their technical and financial support and Dr. Leijun Li, Utah State University, for his technical advice and suggestions.References 1.

for their technical and financial support and Dr. Leijun Li, Utah State University, for his technical advice and suggestions.References 1. Amstutz, H.C., Campbell, P., Kossovsky, N., and Clarke, I.C., 1992, “Mechanism and Clinical Significance of Wear Debris-Induced Osteolysis,” Clinical Orthopaedics & Related Research, 276, pp. 7-18.2. Hamilton, H.W., and Gorczyca, J., 1995, “Low Friction Arthroplasty at 10 to 20 Years: Consequences of Plastic Wear,” Clinical Orthopaedics & Related Research, 318, pp.160-166. 3. Ratner, B.D., Hoffman, A.S., Schoen, F.J., and Lemons, J.E., 1996, Biomaterials Science: An Introduction to Materials in Medicine, Academic Press, San Diego.4. Walker, P.S., and Gold, B.L., 1971, “The Tribology (Friction, Lubrication and Wear) of All- Metal Artificial Hip Joints,” Wear, 17, pp. 285-299.5.

position Rates of Silicon Carbide by Selected Area Laser Depo,sitionY: L. Lee, J V. Tompkins, J M. Sanchez, and H. L. MarcusCenter for Materials Science and Engineering The University ofTexas at AustinAbstractThe deposition rates using pure tetramethylsilane (TMS) as precursor are calculated numerically for a ~od .grown by th~ .Selected Area Laser Dep?sition J?rocess. In particular, the dependence of the kinettcs of deposItion on pressure of TMS IS examIned. The conditions for which volcano d~pos~tion pr?files occur are also investigated. The results show that deposition rate increases wIth IncreasIng pressure and then becomes saturated.

The conditions for which volcano d~pos~tion pr?files occur are also investigated. The results show that deposition rate increases wIth IncreasIng pressure and then becomes saturated. In addition, adsorption-desorption phenomena, rather than effects of reactants depletion, are responsible for the volcano deposition profile observed experimentally.IntroductionSelected Area Laser Deposition (SALD) has seen rapid development during the past few years because of unique characteristics such as its local nature and its potential as a high deposition rate process. Therefore, laser chemical vapor deposition may be used to deposit a wide variety of materials for either prototypes or structural parts.A considerable amount of work has been reported on the deposition of silicon carbide films because of their high mechanical and chemical performances: wear resistance and chemical stability [1]. However, few reports are available in which TMS is used as the precursor gas. Figueras et al.

However, few reports are available in which TMS is used as the precursor gas. Figueras et al. [2, 3, 4] reported several studies that use TMS for conventional CVD systems. They obtained the apparent activation energy, 90 kJ/mol, for the thermal decomposition of TMS using H2 as a carrier gas. Sibieude and Benezech [5] also obtained the apparent activation energy for typical CVD systems using pure TMS precursor. They showed that the activation energy is function ofpressure and substrate temperature. In 1994, Tompkins et al. [6] measured the deposition rate with varying TMS pressure for laser chemical vapor deposition.Volcano-like profiles of deposition were often reported for laser chemical vapor deposition[7]. In general, these deposition profiles can be explained by several factors. They can occur when the reaction rate is much faster than the diffusion rate of the reactants[8].

In general, these deposition profiles can be explained by several factors. They can occur when the reaction rate is much faster than the diffusion rate of the reactants[8]. In this case, the center of the spot is depleted of reactants relative to the periphery giving rise to the volcano-like deposition profile. The temperature dependence of the sticking coefficients also plays a role in the formation of the volcano profiles[9]. Finally, nonlinear kinetics model like Langmuir-Hinshelwood can also explain the volcano profiles [10].The specific problem considered in this paper concerns the growth of a rod by a laser beam using pure TMS as gas precursor. Fig. 1 shows the geometry of the system. Deposition rates as well as temperature and TMS concentration profiles under different conditions are In addition, the predicted deposition rates are compared with experimental results to studied.

Deposition rates as well as temperature and TMS concentration profiles under different conditions are In addition, the predicted deposition rates are compared with experimental results to studied. check the validity of the model.433Mathematical and Numerical ModelSince in SALD the deposition occurs at a locally heated substrate, it is necessary to model first the temperature distribution in the substrate. Several analytical solutions with particular assumptions were obtained using Bessel transforms, Green's functions, or Fourier series expansion technique [11, 12, 13]. However, in this study numerical solutions are obtained by solving the heat conduction equations with variable properties. The governing equation for heat conduction is as follows:where ke(T) and it are the thermal conductivity and the heat source term, respectively.

The governing equation for heat conduction is as follows:where ke(T) and it are the thermal conductivity and the heat source term, respectively. The effective thermal conductivity [14], ke(T) = ks(1- e) / (1 + ~ks / kg), is adopted because silicon carbide powder is used for a substrate to reduce the thermal conductivity. Here, ks and kg are the thermal conductivities for SiC solid and TMS gas, e is the volume fraction of gas in the powder substrate, and ~ is a constant which is about 0.034. The heat source term includes the heat from a laser beam and the heat loss by radiation. The effects of reaction heat and natural convection were assumed to be negligible in this study. A finite volume method was adopted to solve the conduction equations.

The effects of reaction heat and natural convection were assumed to be negligible in this study. A finite volume method was adopted to solve the conduction equations. The calculated temperature profile is then used as a boundary condition when the gas phase is analyzed.The following transport equations, in which the effects of viscous dissipation and thermal diffusion (Soreteffects) are neglected, are solved simultaneously for momentum, heat, and mass transfer:continuity:momentum:energy:species:Here p, f.!, D, C p, and k are respectively the density, the viscosity, the diffusivity, the specific heat at constant pressure, and the thermal conductivity of the gas. Also, mk is the mass fraction of species k, and wis the production rate per unit volume for the kth species. The explanation of other notations in the above equations can be found elsewhere[15].In the above equations, several transport properties appear, which are function of temperature and pressure.

The explanation of other notations in the above equations can be found elsewhere[15].In the above equations, several transport properties appear, which are function of temperature and pressure. Since experimental values of those properties are not available for TMS, the coefficients are derived from standard statistical mechanics theory[16]. The density is determined by the ideal gas law.In addition, chemical reactions have to be taken into account. For most of chemical vapor deposition reactions, only rudimentary models concerning intermediate species are available and almost no kinetic data exist. Therefore, a single overall reaction is used in this434(1)(2)(3)(4)(5)~"'u empirical parameters fitted to measured deposition rate data.investigated by Figueras et al. [4] in the growthdecomposes according to the following overall reaction:on substrate and Neuman condition for the outer boundary.

[4] in the growthdecomposes according to the following overall reaction:on substrate and Neuman condition for the outer boundary. For the energy equations, the the profile in the substrate was obtained from the thermal conduction analysis.applied boundary conditions for the momentum equations areUIV .." " ' ....' ' ' ' equations, a surface reaction rate was used as a boundary condition atsubstrate.TMS is equal to the first order reaction rate.is the binary diffusion coefficient of TMS in CH4, kA = k o/is therate constant, and MTMS is the molecular weight of TMS. The activation ........"'..........'....,.. decomposition of TMS was obtained from the experiment of Sibieude etfor the Also, preexponential factor, ko , in the Arrhenius equation was determined using the no...",... rth rate at torr obtained by Tompkins et al. [6]. The preexponential constant ..........

rth rate at torr obtained by Tompkins et al. [6]. The preexponential constant .......... ,.,u'-'uexperimental growth rate was 2.1 x 108 mls.above conservation equations are discretized by employing a finitecollocated grid. For computation of the pressure field,theDiscussion..... "''V2 shows the calculated properties of TMS using a hard sphere this model, all properties are function ofv~~.. ....,.11 interactions.binary diffusion coefficient which is function of bothandcoefficient is inversely proportional to gas pressure. with radial distance for two different laser powers.powder substrate was 0.2 and the diameter of a laser""''''''' .. ''''' ..... JL'''''' ....' ....

with radial distance for two different laser powers.powder substrate was 0.2 and the diameter of a laser""''''''' .. ''''' ..... JL'''''' ....' .... temperature profiles were close to a Gaussian distribution, t-""......"'....,.""...."'1-·.u·'" occurs at the center of the laser spot.4 shows the typical distribution of temperature for 20 torrC''''1-h''''l'''14l'\C' are almost spherical, it is believed that the diffusion ,",.11"""'V~.:J...... ""...".11""".roOTI''''''',r''Tlr'l.n effects at this low pressure. The TMS concentration UJ.Ul~"'''~''~''Jl'U''''''it is clear thatdiffusion effects prevail over the .........,,,....... ....,..435(6)(7)the decrease in experimental growth rate is due to the contamination of the reactor window by a byproduct of the reaction during the experiments.7 shows the profiles of deposition rate for several different reaction rates.preexponential constant was arbitrarily increased to see whether the volcano profiles occur.

It is clear from Fig. 7 that the depletion of reactants due to high reaction rate indeed causes volcano deposition profiles to form. However, we recalled that the preexponential constant needed to fit the experimental growth rate was 2.1 x 108 mis, for which the volcano effects were not observed. Thus, the effects of reactants depletion do not seem responsible for the volcano profiles observed experimentally.effects of sticking coefficients on the deposition profiles are shownFig. 8.sticking coefficients used in this study is assumed to be 1 if T < TLand 0 for T>TH, varying linearly between the two temperatures. The deposition profiles in Fig. 8 are obtained for TL=1400 with varying TH. The volcano profile starts to form if TH is lower than 1700K. Furthermore, no deposition occurs near the center of a laser spot for TH=1500 K. Thus, the temperature dependence of the sticking coefficients can also explain the volcano profiles.

Furthermore, no deposition occurs near the center of a laser spot for TH=1500 K. Thus, the temperature dependence of the sticking coefficients can also explain the volcano profiles. Since the profiles are strongly dependent upon the two critical temperatures, it will be crucial to use correctandfor numerical modeling.ConclusionThe deposition rate of TMS by Selective Area Laser Deposition has been successfully obtained with a single overall reaction.

The present study leads to the following conclusions: (1) effects of natural convection on deposition rate are negligible since the operating pressure are relatively low; (2) The deposition rate increases with increasing TMS pressure and saturates at high pressure; (3) The temperature dependence of the sticking coefficient or nonlinear kinetics model, rather than the depletion of reactants, seems to be the more likely cause for the volcano like profiles observed experimentally.AcknowledgmentsThis material is based in part upon work supported by the Texas Advanced Technology 1.JI....r"n-.... #JIVV\ under Grant No. CSME-ATPD-157.References1. E. Fitzer and D. Hegen 1979, Angew. Chern. Int. Ed. Ergl., vol. 18, pp. 295. 2. A. Figueras, S. Garelik, and R. Rodroguez-C1emente 1991, " A morphological and Structural Study of SiC layers obtained by LPCVD using Tetramethylsilane", Journal of Crystal Growth, vol. 100, pp. 528.3. A. Figueras, S. Garelik,j. Santiso, and R.

pth of Cut Monitoring for Hybrid Manufacturing Using Acoustic Emission SensorHaythem Gaja*, Frank Liou*+ * Mechanical Engineering Missouri University of Science and Technology, Rolla, MO 65409AbstractLaser Metal Deposition LMD is a hybrid manufacturing process consist of a laser deposition system combined with a 5-axis CNC milling system. During laser deposition many parameters and their interaction affect the dimensional accuracy of the part produced, powder flow rate, laser power and travel speed are some of these parameters. Sensing the acoustic emission during milling marching gives feedback information regarding depth of metal being cut subsequent part dimensions, if an error in dimensions is found certain actions, such as remaching, close loop control, or laser remelting can be carried out to correct it. Thus a reliable hybrid manufacturing management system requires that a depth-of-cut detection system be integrated with the milling machine architecture.

Thus a reliable hybrid manufacturing management system requires that a depth-of-cut detection system be integrated with the milling machine architecture. This work establishes, first a methodology to detect an acoustic emission signal, so that the acoustic emission characteristics of the milling could be analyzed. Second, it sought to relate these acoustic data to machining parameters to detect depth-of-cut.Keywords: Depth of cut Detection, Acoustic emission, artificial neural networkINTRODUCTIONOne of the difficulties in using an adaptive control and tool monitoring system is accurate representation of the variation in machining variables such as cutting speed, feed rate, and depth- of-cut.

In the end-milling process, particular changes in depth-of-cut must be carefully considered to ensure the effectiveness of the control system.Many researchers have sought to control surface errors and radial and axial depth-of-cut using analytical models, simulation, force sensors, and other sensors. Choi [1] suggested an algorithm to estimate the cutting depth based on the pattern of cutting force. He found that the cutting force pattern is more useful for this purpose than its magnitude because its pattern reflects the change in cutting depth. However, magnitude is affected by a number of cutting variables, but not by the depth-of-cut.Yang [2] suggested an analytical method to identify depth-of-cut variations based on cutting force profile features detected during end milling. Based on the profile characteristics of a single-flute, he studied end mill cutting forces and categorized them into three types.

Based on the profile characteristics of a single-flute, he studied end mill cutting forces and categorized them into three types. The same study categorized the cutting forces signals of both the single-flute end mill cutting and the multiple-flute end mill cutting based on the cutting process.Wan [3] predicted the cutting forces and the surface dimensional errors using iteration schemes. Using the finite element method, he devolved a general method to calculate static form errors in peripheral milling of thin-walled structures, and his simulation tool considered the complexity of the workpiece.Li [4] presented a comprehensive time domain model for general end milling processes. The model measures variations in depth-of-cut using mode forms. The model can also consider additional general conditions such as cutting with a large axial depth-of-cut or small discontinued radial depth-of-cut.

The model can also consider additional general conditions such as cutting with a large axial depth-of-cut or small discontinued radial depth-of-cut. In addition to simulating the end milling process this method predicts a number of results for surface profiles and chatter boundaries.Yonggang [5] examined cutting forces and categorized them into six classes according to a combination of cutting depths, and he proposed a finite-element model to study surface dimensional errors in peripheral milling of thin-walled workpieces for aerospace application. Such error prediction keeps the number of surface errors within permissible bounds.To forecast a surface form error with the greatest efficiency and accuracy, Yonggang’s model relies on a set of flexible iterative rules with a double iterative algorithm. Prickett [6] presented an approach that uses ultrasonic sensors for online monitoring of depth-of-cut during the end milling processes.

Prickett [6] presented an approach that uses ultrasonic sensors for online monitoring of depth-of-cut during the end milling processes. The proposed monitoring process tried to contribute to the development of more efficient tool management procedures and supporting infrastructure. However, sensor resolution is an important factor limiting performance.This work used a fitting model with a Neural Network to represent the relationship between the acoustic emission signal and depth-of-cut. The output of the sensor and data of cutting conditions and tool status are fed to a neural network to measure operation quality during machining. After the network was trained, the inference system estimated the depth-of-cut in real time from the experimental sensor signal and the cutting conditions. The results of the monitoring algorithm can warn the operator to take the corrective actions to reach the required depth-of-cut.

The results of the monitoring algorithm can warn the operator to take the corrective actions to reach the required depth-of-cut. The difference between the desired depth-of-cut and the actual depth-of-cut may be a result of incorrect workpiece set-up, tool length offset change (tool wear), or irregularity of workpiece dimensions. Previous manufacturing processes may also lead to errors in depth-of-cut. For example, when a workpiece is manufactured by laser deposition, forging, or casting, the dimensions are not always accurate and uniform.EXPERIMENTAL SETUPFigure 1 shows a schematic diagram of the experimental set-up. The milling process was carried out on a Fadal vertical 5-Axis computer numerical control machine (3016L) using a carbide flat-end mill (0.5 in) to cut deposited stainless steel 316 workpieces. The control interface (National Instrument PXI 7240 and PXI 1250) provided the control and data acquisition.

The control interface (National Instrument PXI 7240 and PXI 1250) provided the control and data acquisition. An acoustic emission sensor (Kistler 8152B211) captured a high-frequency signal. The bandwidth of the AE sensor was 10 to 1000 kHz. The RMS signals were first fed through the data acquisition system and then recoded and processed using Labview software. A 500X digital microscopiccamera was used to detect tool status without disengaging the tool from the tool holder. The tool condition was documented from the bottom edge radius, which was measured in place with the aid of the vision system.Figure 1. Experimental SetupTOOL STATUS CATEGORIZATIONThis study classified tool status according to tool life or tool wear, which is caused by progressive loss of tool material during cutting and which thus changes the shape of the cutting edge. Image J software was used to convert tool wear from a pixel scale to micrometer scale.

Image J software was used to convert tool wear from a pixel scale to micrometer scale. Once the measuring scale was calibrated, tool wear was measured by counting pixels from the vision system and comparing the number with the scale on the reticle Figure 2.The international organization for standardization [7] recommends that the tool be considered worn-out and reached its end point at 300 µm. Here, the output was assigned a value of 1 (for a fresh tool with wear less than 130 µm), 2 (for an average tool between 130 µm and 300 µm), or 3 (for a worn-out tool with wear greater than 300 µm). Figure 3 shows a fresh tool with 10 µm tool wear and Figure 4 shows a worn-out tool with 320 µm tool wear. In both cases, the tool has four flutes with a different level of wear, so the tool wear value represents an average.Figure 2. Tool Wear for 0.5 Inch End Mill425(a)(b)Figure 3.

In both cases, the tool has four flutes with a different level of wear, so the tool wear value represents an average.Figure 2. Tool Wear for 0.5 Inch End Mill425(a)(b)Figure 3. Tool Status (a) The Wear = 10 µm (b) The Wear = 320µmDESIGN OF EXPERIMENTSMost research has focused on the use of a force signal to detect, model, and control radial depth-of-cut and chip thickness [1-5], but this study used an acoustic emission sensor to predict axial depth-of-cut during end milling. The experiments described here were designed to use the most significant factors affecting the acoustic emission signal during the end milling process. Therefore, their outcomes are significant for the computation of depth-of-cut, and the model considers the cutting tool condition and the cutting variables.

Therefore, their outcomes are significant for the computation of depth-of-cut, and the model considers the cutting tool condition and the cutting variables. These factors include depth-of-cut, spindle speed, feed rate, and tool status.A four factor-three level (34) full factorial experimental design with three replications, atotal of 243 cutting tests were run randomly, and a range of cutting conditions were collected.Table 1. Factors and Levels Defined for ExperimentationDepth of cut (mm) 0.5 1 2Cutting Speed (RPM) 1500 3000 5000Feed Rate (mm/min) 40 70 100Tool Status ≤130 µm > 130 µm and ≤ 300 µm >300 µmPREDICTION USING ARTIFICIAL NEURAL NETWORK MODELINGArtificial neural network is statistical machine learning tool established based on the idea of how neurons in human brain work. The neural network consists of layers and nodes called neurons the number of layers and neurons depends on the difficultly of the problem being modeled.

The neural network consists of layers and nodes called neurons the number of layers and neurons depends on the difficultly of the problem being modeled. The input and output layers have neurons equal to the number of the inputs and the outputs respectively. The neurons connected by synapses which take a value from an input neuron and multiply it by specific weight and output the results, neurons have more complicated purpose, they add together all outputs from all synapses and apply activation function.An artificial neural network at least has three layers; input layer, hidden layer, and output layer. If X is the input data vector which in this work is 1 by 4 vector Figure 4, W(1) is weight matrix which is 4 by N matrix , where N is the number of neurons in the hidden layer, and Z(2) is the transfer function of second layer.Figure 4.

Finally activation function is applied to Z(3)in order to obtain the estimate for depth of cut 𝑦′𝑦′ = f(Z(3))(4)Without training the network the estimation error will be very large, training is the process of updating the weight matrix to minimize the cost function 𝐽𝐽 = ∑ 1 2(𝑦 − 𝑦′)2(5)One of training algorithms can be used to train the ANN is a supervised learning algorithm called Backpropagation algorithm which adjust two parameters learning rate and momentum coefficient and keep them between 0 and 1.Equation 5 can be written as𝐽 = ∑ 1 2(𝑦 − f(f(X𝑊(1))𝑊(2))2(6)In order to save time and reduce calculations Gradient Descent method is used to guarantee searching for J in the correct direction and stop searching when smallest J is reached (cost function stops decreasing) by taken the partial derivative of 𝐽 with respect of W ( is positive then the cost function is increasing and vice versa. This method is useful especially in multi- dimensional problems.

This method is useful especially in multi- dimensional problems. Gradient descent can be performed either after using of all training data (batch gradient descent) or after each input–output pair (sequential gradient descent).𝜕𝐽𝜕𝐽), when𝜕𝑊𝜕𝑊The neural network was trained with 147 data points (cutting conditions) to estimate the weights (includes biases) of candidate designs, and 48 data points was used to estimate the non- training performance error of candidate designs and also used to stop training once the non-training validation error estimate stops decreasing. Also 48 data points was used as testing data to obtain an unbiased estimate for the predicted error of unseen non-training data.

Also 48 data points was used as testing data to obtain an unbiased estimate for the predicted error of unseen non-training data. Training, validation, and testing data were randomly chosen from different cutting conditions from the data set consists of 243data points (cutting conditions).Figure 5 illustrates the mean square error versus iteration (Epochs) number, while using the Bayesian regularization training algorithm. 25 neurons was used with in the hidden layer in this work. The network was trained for 90 iterations, at which time the performance was changing dramatically, the best performance was 1.81084 at epoch 65.Figure 5. Neural network performanceESTIMATION OF DEPTH-OF-CUTFigure 6 shows the depth-of-cut estimated at 2 mm, 1 mm, and 0.5 mm with a feed rate of 40 mm/min and a cutting speed of 5000 rpm. Clearly, the system can detect the depth-of-cut with a maximum acceptable error.

Clearly, the system can detect the depth-of-cut with a maximum acceptable error. The accuracy of depth-of-cut estimation depends on the quality of the acquired signal.This work tested the efficiency of the model in estimating depth-of-cut in an interrupted cutting process. As shown in Figure 7, a 25.2 mm slot was made in the workpiece perpendicular to the machining direction. The depth-of-cut was 1 mm, the cutting speed was 4000 rpm, the feed rate was 30 mm/min, and the tool was fresh. 51 second is the time required for the tool to cross the gap (25.2/30), and 25 second is both engagement and disengagement time subtracted from 51 seconds. Figure 8 shows that the model is able to distinguish the slot; thus the system is capable of detecting the engagement and the disengagement of the tool with the workpiece as well as the depth-of-cut.Figure 6.

Shows the depth-of-cut estimated at 2 mm, 1 mm, and 0.5 mm with a feed rate of 40 mm/min and a cutting speed of 5000 rpmFigure 8 shows both the nominal and estimated depth of cut for inclined surface cutting. A 10 mm ramp was created at the end of 60 mm cutting with 2 mm height as shown in the cutting geometry in the figure. The cutting speed was 4000 rpm, the feed rate was 30 mm/min, and the tool was fresh.Figure 7. Interrupted Cutting (a) Nominal Depth-of-cut of 1 mm and Estimated Depth-of-cut (b) Schematic Representation of Cutting Geometry431Figure 8. Inclined Surface Cutting (a) Shows both the Nominal and Estimated Depth-of-cut for Inclined Surface Cutting (b) Schematic Representation of Cutting GeometryAs final test for the efficiency of the depth-of-cut detecting system, a free form surface was made from stainless steel 316 using laser deposition.

The deposited part was first scanned using a 3D scanner, then the part was machined and scanned again as shown in figure 9.Figure 9. Scanned Deposited Material (a) Machined (b) Original (c) RemovedThe difference between the two scans is the machined material. The machined material was sliced to fifty sections as shown in figure 10 and the area of each section was calculated. In order to calculate the depth-of-cut, the area of each section was divided by the tool diameter (12.7 mm).Figure 10. The Difference between the Laser Scans of Deposited Part and Machined Part(a) First Section of Removed Material. (b) Fifteenth Section of Removed Material.Figure 11 shows the measured depth-of-cut from the sections and detected depth-of-cut by the acoustic emission sensor. The feed rate was 60 mm/min, cutting speed 4000 rpm, cutting length about 52 mm and the tool was worn-out. There is some deference between the measured and detected depth-of-cut in several points.

The feed rate was 60 mm/min, cutting speed 4000 rpm, cutting length about 52 mm and the tool was worn-out. There is some deference between the measured and detected depth-of-cut in several points. This error might be caused by the change in the shear strength of the deposited material where the depth-of-cut detection model was made with material deposited at 800 W laser energy and the material tested now was made at 1000 W laser enegy.Figure 11. Measured Depth-of-cut from the Sections and Detected Depth-of-cut for aDeposited Material Detected by the Acoustic Emission SensorCONCLUSIONSThis research investigated experimentally the depth-of-cut and the acoustic emission variations during end-milling of deposited stainless steel 316 with an uncoated tungsten carbide tool under dry conditions, and it studied the correlation between the acoustic emission variation and the depth-of-cut. A full factorial experimental design was used to conduct experiments.

A full factorial experimental design was used to conduct experiments. As a result of this work, neural network model was developed to predict depth-of-cut in end milling.The experimentally determined depth-of-cut values were compared with values predicted by the model, and the model is proved to be capable of predicting depth-of-cut with the acceptable margin of error. The results indicate that this model is robust and accurate. This study provides a depth-of-cut monitoring system for more efficient manufacturing in the future.The Bayesian regularization training algorithm was used to train the network. 25 neurons was used with in the hidden layer in this work, the results obtained after training showed the effectiveness of this approach.The model confirmed the effectiveness of estimating depth-of-cut for inclined surface cutting.

The model is capable to distinguish the slot in an interrupted cutting process; therefore the system is capable of detecting the engagement and the disengagement of the tool with the workpiece as well as the depth-of-cut.The main concern of this work was to detect the depth-of-cut for part made with Laser Metal Deposition, the model showed good agreement between measured depth-of-cut by using a 3D scanner and the predicted depth-of-cut by the artificial neural network model.Future work will investigate signal processing and feature extraction since the root mean square is provided by the coupler and there is no control on low-pass and high-pass filters. A raw signal can be acquired from the sensor, and this signal contains more information than the root mean square signal, which was already processed inside the coupler.

A raw signal can be acquired from the sensor, and this signal contains more information than the root mean square signal, which was already processed inside the coupler. Also, the neural network approach can also be used to estimate feed rate, cutting speed or tool wear when the other cutting parameters are given.References[1]Choi J. G., and Yang, M. Y, "In-Process Prediction of Cutting Depths in End Milling," Int. J. Mach. Tools Manuf., 39(5), pp. 705–721. [Inspec] [ISI] , 1999[2]Yang Liuqing, DeVor Richard, Kapoor Shiv, “Analysis of Force Shape Characteristics and Detection of Depth-of-Cut Variations in End Milling,” J. Manuf. Sci. Eng. 127, 454 , DOI:10.1115/1.1947207, 2005.[3] Wan M., W.H. Zhang, “Efficient algorithms for calculations of static form errors inperipheral milling,” Journal of Materials Processing Technology 171, 156–165, 2006.[4]Li, H., and Shin, Y.

This paper will outline new approaches to the design and manufacture of personalised foot and ankle-foot orthoses (FO and AFO) using additive fabrication technology. The research is addressing the need for specific software design tools for orthosis design which enable their properties to be locally tailored within a mass customisation framework. Structure/material testing to support that activity is also being undertaken and will be described.IntroductionAn orthotic device is often prescribed to support, re-align or redistribute pressure across part of a person’s musculoskeletal system. The use of these devices can lead to a reduction in symptoms, improvement in function, and may result in an increase of the patient’s overall quality of life.

The use of these devices can lead to a reduction in symptoms, improvement in function, and may result in an increase of the patient’s overall quality of life. Customised devices have been found to be most effective for the treatment of many conditions in terms of clinically relevant benefits and patient compliance, most likely a result of taking the individual patient’s anatomy and functional requirements into account during the design process (Hawke et al., 2008, Trotter & Pierrynowski, 2008). However, the production of these customised orthotic and prosthetic devices is currently based around artisan skills.

However, the production of these customised orthotic and prosthetic devices is currently based around artisan skills. The processes involved are time-consuming and require experienced craftpersons who generally make their decisions based on experience and trial and error, rather than systematic engineering and evidence-based principles.Furthermore, the mechanical properties of these traditionally produced orthotics can only be grossly estimated since factors such as controlling wall thickness can only be partially achieved. Overall this approach can result in inconsistent designs and patient care.The work reported in this paper examines the feasibility of using an additive fabrication (AF) technique for the manufacture of customised orthoses and prosthetics and describes the potential of this approach for improving lead times, quality, consistency and patient care.

The design freedom made available though this method presents a range of opportunities that are not possible with current orthotic CAD/CAM packages. Examples of where improvements to current systems can be made with tailored software solutions. Some examples of manufacturing prosthetic devices are also mentioned in this paper. This is relevant as the current manufacturing of both prosthetic and orthotic devices is very similar and the research done so far regarding these devices has also been similar in nature.Additive ManufacturingFigure 1 gives a schematic overview of the traditional versus an AF process.

Additive fabrication processes have been available to produce low volumes of components with short lead times since the early 1980’s (Levy et al., 2003), and have more recently been exploited in the production of medical devices, including the manufacture of customised in-the-ear hearing aid shells (Tognola et al., 2003), and the creation of drill guides for dental surgery (Kim et al., 2008), and has also been evaluated for their potential in creating ankle-foot orthoses (AFOs) (Faustini et al., 2008).Figure 1. The traditional orthoses manufacturing process (left) and the same process using additive fabrication technology (right).It is suggested that a mass customisation system to deliver orthoses, together with exploitation of the design freedom offered by the manufacturing method, will give the devices produced through the AF approach significant clinically relevant potential.

Some of the advantages of this approach are as follows:Faster production over the traditional process and more consistent quality.A more comfortable experience for the patient during the prescription process (no plaster casting required).

Optical 3D scanners can be used to capture the patient geometry in a precise and consistent way.Fewer and less experienced technicians can be used as there is a reduced need for manual work and hands on experience is not as crucial.Design rules and company specific protocols, lending themselves more easily to systematic evaluation of finished products, can be implemented into the design system.The orthoses/prosthesis designs can be archived and reproduced when needed.A reduction in the production management required as a result of a more streamlined production processes.Fewer items of production equipment and less storage space needed.More possibilities in new product development.FeasibilityTo establish the feasibility of AF for orthotic and prosthetic manufacturing, several research efforts have been made in recent years:In research conducted at the University of Texas, Austin, the feasibility of using AF for prosthetic socket manufacturing has been investigated in a number of projects.

In one thesis project, compliant structures were developed and analysed using the finite element method (FEM). In a patient trial, it was found that contact pressures between the residual limb and the AF produced socket could be significantly reduced with an integrated compliant surface (Faustini 2004, Rogers et al., 2008).Further research was conducted in Austin to investigate the feasibility of custom made AFO’s and on how to adjust their stiffness. It was concluded that the AF approach is well suited for AFO production (Faustini et al., 2008).In research by the author, it was concluded that the clinical performance of foot orthoses (FOs) fabricated using selective laser sintering (SLS) was comparable to those produced using traditional methods in a seven patient trial where the patients were assessed using gait analysis and questionnaires (Pallari 2008, Pallari et al., 2010).Other, non-technical aspects of the “shift” to AF have also been investigated.

Wagner investigated how the role of prosthetists and prosthetic workshop technicians would change if the manufacturing of prosthetics would change to utilise AF (Wagner et al., 2008). They concluded that the more experienced prosthetists would have to learn new computer skills, but they would still utilise their experience and craft skills. The more junior technician’s role would be reduced as the need for their craft skills would be reduced.A literature survey by Rogers et al reviewed ten research projects investigating the feasibility of AF orthotics or prosthetics (Rogers et al., 2007).

Some limitations to the AF approach are identified (cost and lack of suitable design software) but it is concluded that the basic concept of using AF for orthotic and prosthetic production is viable.A common feature of all the studies which involved working with patients was that the number of patients in each study was small (as would be expected for an initial study) and the time to customise the devices and the various integrated features was not seen as an issue. However, in orthotic practice, time is however not always abundant and the interventions tothe positive casts need to be made quickly. With AF, the final orthotic product can be designed directly, but the same time pressure problem remains.

With AF, the final orthotic product can be designed directly, but the same time pressure problem remains. Adding various design features further increases this design time.Optimising the properties of orthoses with finite element analysisDue to the small number of manufacturing constraints, AF and the SLS process in particular can be efficiently used to fabricate tailored, patient-specific orthotics with a predefined optimal stiffness level. The finite element method is potentially a useful tool here and has been used to assess the mechanical characteristics and functional performance of AFOs (Chu, 2001, Syngellakis et al., 2000).Following traditional methods, a custom-made polypropylene AFO was designed for a healthy test subject (female, 25 years old). The model was scanned and converted into a surface shell (3-matic 5.01, Materialise, Haasrode, Belgium) based on which a finite element mesh was generated (Patran 2008r1, MSC Software, Gouda, the Netherlands).

Boundary conditions simulating dorsiflexion movement were implemented in the numerical model (Creylman et al., 2009). The effect of different materials and different design characteristics on functional parameters of ankle foot orthosis was studied (Muraru et al., 2010). These results combined with features of structural optimisation could lead to optimal and personalized designs of such devices.Topology optimization was used to find the optimal material distribution for the AFO described above and presented in figures 2 and 3.Figure 2: A 3D model of an ankle foot orthosis is presented in A. In B the inner surface is separated and in C the initial design space of 61920 square elements (global element size = 1 mm) is shown.Figure 3.

In B the inner surface is separated and in C the initial design space of 61920 square elements (global element size = 1 mm) is shown.Figure 3. In A the initial design space is presented, in B the optimisation solution and in C the .stl description of the optimised solution.The initial design space consisted of a shell of 4-noded quadrilateral elements which described the inner surface of the physical orthotic (figure 2C).The density method (Bendsoe & Sigmund, 1999) implemented in MD Nastran (MD Nastran R3b, MSC Software, Gouda, The Netherlands) was used for topology optimisation. The design variables in the optimisation problem were the normalized densities of each element from the design space. A power law is used to relate the density with the material properties:Where ρ0 and E0 are the fully solid density and Young’s modulus respectively. A penalty factor p is introduced to enforce the design variable to be close to a 0-1 solution when p>1.

A penalty factor p is introduced to enforce the design variable to be close to a 0-1 solution when p>1. Inthis way, intermediate densities are penalized and the result of the optimisation is a black and white structure (with clearly defined solid and void regions) (MSC Software 2007).The load implemented in the analysis model was static equivalent to dorsiflexion movement. In the optimisation process this is considered independent and but in relation to the mesh. The objective was to minimize the compliance while the mass target was 50%.The final, optimised AFO shape is presented in figure 4. The triangulated surface generated by the analysis was imported into 3-matic 5.01 and aligned with the original scanned AFO. The curves separating the open and closed areas were then used to cut material from the AFO model. Alternatively, only some material could have been removed from the model in indicated locations. These types of CAD operations are easy to automate.Figure 4.

Alternatively, only some material could have been removed from the model in indicated locations. These types of CAD operations are easy to automate.Figure 4. A custom made AFO (A) with material cut off according to the optimised solution. Anterior view in B and the posterior in C.New approaches to the design and manufacture of foot and ankle-foot orthosesReproducing the performance of current state-of-the-art orthotics is an important step in investigating the viability of AF methods and prior to bringing them to the market. Beyond this however, there is potential to integrate functional features in the orthosis and to optimise their properties in ways not possible with traditional methods.Some of these possibilities have been described in the literature, and further examples are presented here for FOs (Faustini 2004, Rogers et al., 2007, Faustini et al., 2008).

For example, patterns of conforming beams, such as the ones presented in figure 5, can be integrated on the top surface of the orthoses. These beams will deform when placed under aload. They could perhaps be added to areas of the orthotic surface based on interface pressure measurements or visual observation of irritated areas on the foot surface. When combined with a thin cushioning layer which is traditionally adhered to the upper surface of the orthosis, they can be used to provide different cushioning properties for different areas of the foot.Figure 5. A conceptual drawing of different beam structures that can be integrated on the surface of an orthoses.The beams could be linked to a pressure measurement as follows. The pressure measurement output, usually visualised in different colours and contours for the regions where different pressures were measured, as shown in figure 6.Figure 6.

The pressure measurement output, usually visualised in different colours and contours for the regions where different pressures were measured, as shown in figure 6.Figure 6. Different pressure regions from a pressure plate measurement projected onto the foot (Alba Chiropractic 2010).This measurement data (specific surfaces or contour curves) can then be exported as greyscale bitmap images. These image files could then be imported into the orthotic CAD system and projected onto the relevant surface on the orthoses. Via preset threshold values, the cushioning beams could then be designed and patterned along the specified high pressure areas.Other examples of functionality that have been built into foot orthoses (Pallari 2008):Rearfoot wedges. The function of a rearfoot wedge is to realign the foot during the gait cycle by rotating the rearfoot relative to the rest of the foot.

The function of a rearfoot wedge is to realign the foot during the gait cycle by rotating the rearfoot relative to the rest of the foot. The front part of the orthoses can be kept flat, while the heel part can be tilted and a rigid heel part designed accordingly. The “blending” of the frontfoot and the rearfoot parts of the orthoses can be seen in figure 7.ABFigure 7. Foot orthoses with an integrated rearfoot wedge. Medial view in A and anterior view in B.Metatarsal pads, bars and cut-outs. These come in a variety of shapes and sizes and are used to redistribute the plantar pressure in the forefoot.Heel and arch supports. These define a “cup” in which the heel will be positioned and the arch support which normally prevents the foot arch from collapsing.Sensor integrationA further opportunity resulting from the design freedom provided from AF is the potential to embed sensors into an orthosis.

These could be used to collect data on a wide range of normal, day-to-day activities over an extended period of time. Data recorded by the embedded devices could provide valuable feedback to the patient or prescribing clinician if the desired treatment/rehabilitation effect intended by the clinician when prescribing a certain orthotic device is being achieved.From a recent literature review on in-shoe sensing technologies covering pressure, temperature, shear, humidity, as well as additional biofeedback indicators the following key specifications have been observed:The measurements taken should be sufficiently accurate and repeatable under a large number of loading cycles and in the different environments (e.g. changing temperatures and humidity levels) that the orthotic will experience in-shoe.They should not affect the performance of the orthoses or cause any discomfort to the patient.The sensor should be reliable.

 Data should be stored with a high degree of fidelity and be easy to access, either through the implementation of wireless technology or with built in data loggers attached to the sensor. Different attempts have been proposed to embed devices for biofeedback indicators into insoles, orthotics or even footwear, for example plantar pressure (Pataky et al., 2000), humidity (Morley et al., 2001) and plantar blood flow detectors (Cobb & Claremont, 2001). Rutkove et al. in 2007 for instance used the iButton, a low cost temperature sensor with built in data logger, to carry out a study comparing temperatures measured in-shoe on the dorsal surface of the foot with ambient levels over the course of two days. Activity was also monitored using accelerometers and a diary was kept by the subjects. The sensor was found to be robust and reliable and the authors recommended its use for clinical and research purposes.

The sensor was found to be robust and reliable and the authors recommended its use for clinical and research purposes. However, to date few sensing technologies are being applied to orthotics.There is typically a lack of studies which describe long term continuous measurement using in-shoe devices. Researchers have attempted to measure a wide range of physiological factors regarding for instance the diabetic foot with generally high levels of success, however the majority of experiments were either non-ambulatory or set in a laboratory with only a relatively small number of steps being measured, with even the longest studies only lasting for a few hours.Biofeedback type systems aimed at causing the patient to adjust their gait have been proven to be effective in a number of studies, however again these have generally only been over short time periods and with low numbers of subjects.

There remains a great deal of potential in this area for more innovative uses of different types of feedback mechanisms, the range of physiological data collected, and the exploitation of the data.AF technologies can enable the sensors to be placed very precisely in the correct place and an unobtrusive way. To do this in an operational setting, the placement has to be fast, easy and all the necessary wires etc need to be accommodated in the design. This is not currently feasible in any orthotics design system and would require a custom application to be developed.Software design tools for orthosis designSeveral different ways of utilising the design freedom of AF in FOs and AFOs have been described. To make these new types of devices in a production operation, suitable CAD systems are required.

To make these new types of devices in a production operation, suitable CAD systems are required. Existing orthotic CAD/CAM systems are not ideal for these new technologies because:The interventions enabled by the software focus in manipulating the positive cast (a duplicate of the patient’s foot or leg), not in creating the final orthotic product.The file input options are usually restricted to photographs and pressure measurement data. FE data for example cannot usually be imported.The file output options are usually restricted to those compatible with milling machines.AF is not considered as a manufacturing option for the end product or of the positive and complex surface manipulation operations are not possible.Automation possibilities, enabled by macro’s for example are not there.Generally speaking, only FO CAD systems allow the design of the end product which is then milled from foam blocks.

To make AF a viable, commercial option for orthotics manufacturing, software solutions are needed to utilise AF to its full potential.Rogers et al. in 2007 summarise the same conclusion as follows: “While SFF will not replace conventional fabrication; it can be another tool in the prosthetist’s toolbox” and further “The cost of the software is the price of entry. The lack of availability of such software is the barrier.”ConclusionsThe latest literature indicates that the basic assumption of using AF to make FOs and AFOs is a feasible one. Further research illustrates how the shape of the orthotic device can be altered to save weight, to have more functional properties, and can be fitted with external sensors.While orthoses can be engineered to a very high degree of user specific customisation through the incorporation of gait and surface pressure measurement analysis into the design process, this is not done in current clinical practice.

This is mostly because of time, cost and manufacturing constraints. The orthotic and prosthetic industry does not have a tradition of engineering and expert designer familiar with finite element analysis for example are not common.

The orthotic and prosthetic industry does not have a tradition of engineering and expert designer familiar with finite element analysis for example are not common. Where the design work is done at a case-by-case basis, using current engineering software, the time it takes to make one design may increase the cost per product significantly.These issues can be partially solved by developing specific software packages to enable the full potential of AF to be utilised in orthotic and prosthetic product design and in creating completely new kinds of products, changing the industry currently restricted by old and inefficient manufacturing methods.AcknowledgementsThis work was funded through the European Commission Framework Seven Program (grant number NMP2-SE-2009-228893) as part of the A-Footprint project (www.afootrint.eu).ReferencesAlba Chiropractic, (http://www.albachiropractic.co.uk/gait_scan.html), 12:00, 7.8.2010.Bendsoe M.P. and Sigmund O.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDesign and Additive Manufacturing of a Composite Crossflow Heat ExchangerT. Mulholland, R. Felber, N. Rudolph Department of Mechanical Engineering, University of Wisconsin, Madison, WI 53706AbstractAdditive manufacturing (AM) with composite materials reveals new possibilities for direct manufacturing of end-use products, breaking the paradigm of 3D printing as only a prototyping or pre-production technique that has been the norm for many AM technologies. A crossflow air-to-water heat exchanger (HX) was designed for manufacturing via fused filament fabrication (FFF). Design iterations improved the manufacturability, considering issues such as geometric fidelity, watertightness, print time, support material, and manufacturing cost.

Design iterations improved the manufacturability, considering issues such as geometric fidelity, watertightness, print time, support material, and manufacturing cost. Carbon fiber fillers enhanced the thermal conductivity of the base polyamide resin, allowing for thermal HX performance comparable to conventional aluminum finned tube heat exchangers. The anisotropic thermal conductivity impacts the heat exchanger performance. The design and manufacturing challenges reveal additional routes to continued performance gains as the HX is scaled up to an 8 kilowatt product.Keywords: 3D Printing, FFF, Thermal Conductivity, Additive ManufacturingIntroductionMaterial extrusion-based additive manufacturing (AM) technology, such as fused filament fabrication (FFF), is possibly the single largest 3D printing process by number of printers, especially considering the availability of these devices as inexpensive, desktop printers that frequently appear in schools and makerspaces.

However, FFF is still widely considered a prototyping technology, and has not yielded many prominent examples of manufactured products, in contrast to AM processes that are used for making shoe soles [1] or printed bones [2].However, FFF has some unique advantages over other AM technologies that can be exploited to make economically viable manufactured parts. FFF uses low-cost material compared to other technologies. Powder materials can cost $175 to $250 per kg, compared to traditional injection molding materials that cost $2 to $3 per kg [3]. FFF materials, however, cost between $30 and $100 per kg, with the higher price reflecting specialty materials that do not benefit from economies of scale. Moreover, FFF materials will cost much less when production is ramped up, since the production method is the well-known, low-cost plastic extrusion technology. In contrast, powder materials are made through more expensive processes: cryogenic grinding, for example.

In contrast, powder materials are made through more expensive processes: cryogenic grinding, for example. Because these FFF materials are produced by extrusion, they can also contain high loads of fillers. Powder materials are not suitable for high filler contents, since the separate polymer and filler particles can segregate during powder application. If the filler is incorporated in the powder, it must be extremely short, and the reinforcing or conducting effects of the fibers are minimized. However, FFF materials have been loaded to 40 %vol or higher [4]. This high filler content can allow substantial enhancement of different properties, depending on the filler material, such as strength, stiffness, thermal conductivity, or electrical conductivity.2641Finally, FFF machines are substantially less expensive than some other technologies, particularly so for powder-based technologies.

While an FFF machine can range from $2500 to $40,000 for a high quality desktop machine up to a mid-range professional machine [5], SLS machines cost $100,000 and upwards [6]. Therefore, if FFF technology is matched with the right combination of material and product, there is a potential for low-cost manufacturing to be viable.Electrically conductive FFF materials can be made by adding conductive fillers to the base polymer resin, and these have been available for some time. The thermal conductivity can also be enhanced through the addition of fillers, although this generally requires higher filler loads [7] and benefits especially from fillers with a high aspect ratio, such as fibers [8]. It is understood that the fibers will tend to align in the flow direction in injection molding [9], which leads to anisotropic thermal conductivity with the highest conductivities associated with the major flow direction. The same behavior can be expected in the flow direction in FFF processing.

The same behavior can be expected in the flow direction in FFF processing. These anisotropic properties can influence the design or manufacturing of a part, depending on the application requirements. Thermally conductive polymer composites have been used in injection molding for heat transfer applications, such as heat sinks [8]. However, AM allows parts to be produced with a high geometric complexity without additional costs, and can thus allow some atypical geometries to be manufactured economically. This work focuses on FFF printing of an air-cooled cross-flow heat exchanger, which enhances the heat transfer rate compared to traditional corrugated aluminum heat exchangers by raising the convection coefficient using many pin fins. However, printing these complex geometries yields some manufacturing challenges.

However, printing these complex geometries yields some manufacturing challenges. Design requirements, manufacturing restrictions, and implications of the anisotropic properties should all be considered for the manufacture of a thermoplastic composite air-cooled heat exchanger.MethodsMaterial A commercially available polyamide material with carbon fiber fillers, called Onyx (Markforged), was used for the print tests in the study. Burnoff tests showed that the material is approximately 14 %wt carbon fiber. The material was characterized by laser flash analysis, to measure the thermal diffusivity, and differential scanning calorimetry, to measure the specific heat capacity, but these measurements are not discussed here; more information is available in [10].

However, the flow-direction thermal conductivity of the Onyx material at room temperature was calculated from these measurements to be approximately 0.9 W/m-K, while the perpendicular conductivity is 0.3 W/m-K, similar to the thermal conductivity of an unfilled polyamide. This anisotropic conductivity has implications that are discussed later.A Makergear M2 printer, version E was used for printing samples. This FFF printer uses a direct-drive extruder mounted on a gantry with an axis in the x-direction, and the build plate is mobile in the z- and y-directions. A 0.35 mm stainless steel nozzle was chosen for the prints to resist the abrasion caused by the carbon fibers.

A 0.35 mm stainless steel nozzle was chosen for the prints to resist the abrasion caused by the carbon fibers. Since polyamide is hydrophilic, and a high water content can cause printing defects, the material was stored in a dry box with a tube connecting to the printer, allowing the material to be kept dry until printing.2642Micro-Computed Tomography (µCT) Some printed samples were scanned using a Zeiss Metrotom 800 µCT instrument. The technique takes many X-ray images of a sample as it is rotated on a stage; the reconstruction algorithms can convert the many 2D images into a 3D dataset. Scans of small and large printed parts had a resolution in the range of 20 to 40 µm, depending on the size of the object.DiscussionThe heat exchanger (HX) design is shown below in Figure 1.

Scans of small and large printed parts had a resolution in the range of 20 to 40 µm, depending on the size of the object.DiscussionThe heat exchanger (HX) design is shown below in Figure 1. Water enters from one side, flows through half of the internal water channels to the opposite side, where it is recombined in a manifold, then travels through the remaining water channels to reach the outlet. At the same time, air flows perpendicular to the water, through air channels that contain pin fins to enhance the convection coefficient. On the right in Figure 1, the arrangement of airfoil pins through the air channels is shown. Earlier designs began with circular pins, since correlations for flow velocities and heat transfer coefficients are well-known, facilitating the design and verification of the HX. These circular pins very well illustrate some of the problems with using FFF to produce small geometries with sizes on the order of the nozzle diameter, as discussed below.Figure 1.

These circular pins very well illustrate some of the problems with using FFF to produce small geometries with sizes on the order of the nozzle diameter, as discussed below.Figure 1. Left: the cross-flow HX with pins and internal water channels. Right: the cross-section of the pins is an airfoil.To understand the issues with printing small features, it is useful to consider a horizontal cylinder with a diameter equal to the nozzle diameter of the printer: 0.35 mm in this example. When the 3D model is converted into a STereoLithography (STL) file, the surface of the cylinder is tessellated with triangles, and the circle is represented with a series of line segments. This approximation of the circle as a polygon can cause a small loss in the represented circle diameter, which can in turn cause the toolpathing software, known as a slicer, to ignore the feature.

This approximation of the circle as a polygon can cause a small loss in the represented circle diameter, which can in turn cause the toolpathing software, known as a slicer, to ignore the feature. To put it another way, some toolpath generation software, such as the free, open-source Slic3r program, will sometimes ignore features that are smaller than a nozzle diameter. Thus, the original 3D model must be slightly enlarged, to account for the loss in diameter during the STL conversion. This is illustrated in Figure 2 below.As the slicing program creates the 2D layers from the STL file, based on the user-defined layer height (0.10 mm in the example in Figure 2), it will encounter layers containing sections of the cylinder that are thinner than the nozzle diameter. These are ignored, so depending on the quality of the STL file, there may be only one layer that leads to a bead to be printed, illustrated on the right in Figure 2.

These are ignored, so depending on the quality of the STL file, there may be only one layer that leads to a bead to be printed, illustrated on the right in Figure 2. If this bead is a bridge, like the pin in the example, then the resulting bead will be a small stretched ellipse: quite different from the original, designed cylindrical pin.2643Figure 2. Illustration of geometric losses. Left to right: an axial view of a cylinder designed larger than intended, theSTL conversion reduces the circle width, the slicing software divides the shape vertically by layer height, and the only sliced layer wide enough is printed with an elliptical shape.The pins printed in early versions of the HX were cylinders with a diameter of 1 mm. These cylinders bridge a relatively small gap, between 2 and 3 mm, which is much smaller than the maximum bridge length.

These cylinders bridge a relatively small gap, between 2 and 3 mm, which is much smaller than the maximum bridge length. Figure 3 shows the maximum packing of beads into the space for the given pin diameter of 1 mm, nozzle diameter of 0.35 mm, and layer height of 0.2 mm. However, if the intended geometry (dashed line) were shifted up slightly with respect to the print layer (i.e. the pin is slightly farther from the build plate), then only seven beads would fit in the space. And as is evident in the µCT image, shown on the right, there is an interaction between the STL file and the toolpath generation that led to only six beads printing in most cases. Regardless, these circular pins do not always consolidate well, leaving holes on the bottom at times, which can lead to disruptions in the airflow, raising the pressure drop in the HX.Figure 3. Left: a circular pin as it would be printed with a 0.35 mm nozzle and 0.2 mm layer height.

Left: a circular pin as it would be printed with a 0.35 mm nozzle and 0.2 mm layer height. Right: actualcircular pins printed with those settings and scanned with µCT.In order to move to using more complex pin shapes, which were promising for improving the heat transfer per mass and heat transfer per pressure drop ratios, a test piece was designed, which is described in detail in other work [11], as part of this project. The part includes five pin shapes at three different sizes each, as pictured below in Figure 4. The part was printed and scanned via µCT, which allowed some analysis of the geometric fidelity. An example intended cross section of the pins is pictured on the left in Figure 5, whereas the actual print cross section is given on the right; the white color is the material, and the black color is air.

Clearly, there are many air gaps in the part, in some cases due to the toolpath (for example, the turns where the pins meet the wall) and in others due to the cross section; where the width of the pin is an uneven multiple of the nozzle diameter, the slicing program ignores the extra space. It is worth noting that these spaces could theoretically be filled, but there is no commercial or open-source slicer known to the author that will do that.2644Figure 4. A test piece was designed [11] to test geometric reproducibility of several fin shapes.The airfoil-shaped pins were reproduced fairly well, with few noticeable voids. These are shown in the bottom row of the µCT scan in Figure 5, and a perpendicular cross-section is shown in the bottom row of Figure 6. Although there are some unavoidable voids in the pins, the overall airfoil shape is well-defined.

Although there are some unavoidable voids in the pins, the overall airfoil shape is well-defined. Other shapes, such as the ellipsoids, showed flattening on the top or bottom due to the toolpaths, or poor consolidation, for example in the case of the circular pins.Figure 5. Left: intended xy-plane cross section of the different pin geometries. Right: actual printed geometry showingvoids. White represents material and black represents air.Toolpath issues also affected the walls between the air and water channels, which directly led to a loss of watertightness. Figure 7 shows one of the problems. The left illustration shows toolpaths in red and blue. Examining the blue path, it is seen that the printer produces this geometry by moving in an H-like pattern across the channel, producing the walls and pins simultaneously. In the center image, a closeup view of one pin shows the detailed toolpath at this point. First, the outer edges of the pin are printed.

In the center image, a closeup view of one pin shows the detailed toolpath at this point. First, the outer edges of the pin are printed. Later, the center of the pin is printed, and a rotational motion attempts to fill the space at the intersection of the pin and wall. This motion reheats the frozen wall and causes deformations, which can lead to holes in the wall. These are shown in Figure 8 on the left.2645Figure 6. Designed (left) and printed (right) elliptical (top) airfoil (bottom) pins as seen in CT scans.The right image in Figure 7 shows a custom toolpath, which had to be created by combining separate pin and wall STL files, then assigning a different FFF extruder to each STL. The slicer can be told that the printer has two extruders that occupy the same coordinates, and the tool change-related G-code commands can later be removed. Now, the pin is printed perpendicular to and on top of the previously printed wall, layer by layer.

Now, the pin is printed perpendicular to and on top of the previously printed wall, layer by layer. This eliminates the rotational motions at the wall and improves the wall quality, as seen in Figure 8 on the right. Unfortunately, this degree of toolpath control is not currently possible on arbitrary geometries, and there are still imperfect moves in the improved toolpath. For example, the right side of the pin is first printed in space (Figure 7, right), instead of beginning at the wall where it could be anchored.Figure 7. Left: an overview of the pin and wall dividing the air and water channels. Center: inset from left, theunintuitive automatically generated toolpath. Right: a customized, improved toolpath.2646Figure 8. Marks from pin printing can be seen through the wall. Left: The auto-generated slicing path leads to holes inthe walls, compromising the watertightness.

Marks from pin printing can be seen through the wall. Left: The auto-generated slicing path leads to holes inthe walls, compromising the watertightness. Right: a custom slicing path improves the wall quality.When considering the thin walls between the air and water channels, it is also instructive to consider the anisotropic thermal conductivity of a fiber-filled material. Recalling that the thermal conductivity is high in the fiber-aligned direction, which is the same as the flow direction, then the fibers should be aligned along the toolpath in an FFF print.This presents a fundamental challenge in HX printing; the low conductivity of plastics (even filled plastics) drives the HX design to smaller and smaller features to shorten conductive paths, but small features greatly constrain the possible tool paths. The two most important features in the design are the fins and the walls dividing the two fluids.

The two most important features in the design are the fins and the walls dividing the two fluids. The fins will ideally conduct heat from the fluid to the walls (in-plane) at a very high rate. However, the walls will ideally conduct heat from one fluid channel into the other (through-plane), but not conduct along the length or width of the wall (in-plane). As illustrated in Figure 9 below, the fibers are oriented in the ideal direction along the axis of the pin, but they are oriented perpendicular to the ideal direction in the walls.Figure 9. Schematic drawing of the print direction in the walls and fins.For a given thermal conductivity, the HX wall should be as thin as possible. However, a very thin wall means that the toolpath during printing will be oriented along the length of the wall, putting the fibers in the worst orientation. One alternative could be to print the walls with a2647perpendicular path, creating them with S-shapes.

One alternative could be to print the walls with a2647perpendicular path, creating them with S-shapes. This requires thicker walls, which raise the weight of the heat exchanger, so the effect on the overall heat transfer per mass performance is unclear. A better solution might be to use an isotropic material in the walls, and only deposit the fiber-filled material in the pins. Dual material printing like that is greatly facilitated by FFF compared to other technologies. Regardless, it is clear that the thermal anisotropy will affect the heat exchanger design and performance.ConclusionAlthough FFF printers and slicing software continue to advance, there are real manufacturing challenges associated with small features and watertight structures. There is some loss in geometric fidelity during the conversion of solid 3D model to STL, and there is a larger loss when filling a relatively small shape with the elliptical beads associated with FFF.

Slicing software does not allow fine control over the toolpath in different areas of a part, so currently, users are stuck with workarounds and tricks to address those problems. Nevertheless, advances have been made improving the quality of thin walls with these pin structures. Finally, although these other issues focus on the small-scale structure of the part, the microstructure also matters; that is, the fiber orientation within the material has an important influence on the part performance in this heat transfer application. This must be considered when designing the part, simulating its performance, and developing a manufacturing scheme for enhanced performance.Future WorkThese manufacturing issues were presented in the context of an air-cooled heat exchanger, but they are relevant for technical parts with a small feature size in general.

The predictability of the differences between the designed part and the printed geometry is important, especially considering the functional aspect of the small features, cooling fins in this case; this predictability will probably improve as computer-aided design (CAD) software continues integrating printer control schemes, which is beginning to appear in some products such as SolidWorks and AutoCAD. This integration will hopefully lead to some degree of control or influence over the toolpath, since it is easiest to select features before the model is converted to STL.

This integration will hopefully lead to some degree of control or influence over the toolpath, since it is easiest to select features before the model is converted to STL. Desired features would include a print order (walls before pins) to improve watertightness, a preferred print direction (print pins moving in the y-direction before moving in the x-direction) to mitigate stringing or reduce the number of accelerations and the print time, and the ability to set different print parameters in different regions of a part.At the same time, this project will continue to focus on materials and application development. Armed with an understanding of the design issues, materials will be further developed to raise the thermal conductivity, in the flow direction but especially perpendicular to the flow direction. Many filler materials are potentially viable, from aluminum and copper flakes to carbon fiber and graphite.

Many filler materials are potentially viable, from aluminum and copper flakes to carbon fiber and graphite. The relationship between the filler shape and the maximum filler content, limited by nozzle clogging, must be better understood. Variety and predictability in the materials will lead to better product design, which will lead to further examples of FFF- manufactured parts.2648AcknowledgementsThe authors would like to thank the U.S. Department of Energy ARPA-E for funding.References[1] Adidas,“AdidasFuturecraft,”2017.[Online].Available:[2]http://www.adidas.com/us/futurecraft. [Accessed: 2-Aug-2017]. “3D-printed skull implanted in patient,” UMC Utrecht, 2014. [Online]. Available: http://www.umcutrecht.nl/en/Research/News/3D-printed-skull-implanted-in-patient. [Accessed: 2-Aug-2017].[3] M. Cotteleer, “3D opportunity for production: Additive manufacturing makes its(business) case | Deloitte University Press,” 2014.[4] M. Nikzad, S. H. Masood, and I.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesign and additive manufacturing of a patient specific polymer thumb splint conceptMazher Iqbal Mohammed1 and Pearse Fay21Deakin University, School of Engineering, 75 Pigdons Road, Waurn Ponds, Geelong, VIC 3216, Australia2Deakin University, School of Health & Social Development, 1 Gheringhap St, Waterfront Campus, Geelong, VIC 3220, AustraliaAbstractTraditionally, upper limb splints often fall short of being optimal with respect fit and patient expectations, resulting in a lack of use and no treatment of the underlying condition.

In this study we address several current limitations and examine the feasibility of using 3D optical scanning, Computer Aided Design (CAD) and low cost 3D printing as a tool to create more ergonomic and efficacious splints for patients suffering from compromised musculature or trauma of the thumb. Optical scanning allows for a non-invasive and rapid means to reproduce the surface topology of a person’s hand and this data was used as the template for the device design. We explore the use of CAD to create a more aesthetically pleasing and functional splint, enhancing both comfort and potential moisture release.

We explore the use of CAD to create a more aesthetically pleasing and functional splint, enhancing both comfort and potential moisture release. Finally, we demonstrate that low cost polymer printing can allow for rapid design evaluation and production of a final, usable device.Keywords: Splint, 3D Printing, CAD, Patient Specific, FFF, Polymer, OrthoticCorresponding Author: mazher.mohammed@deakin.edu.au1 IntroductionThe use of splint devices is the traditional benchmark treatment option for patients suffering from a large range of musculoskeletal conditions of the upper limb, both acute and chronic [1, 2]. Health professionals, such as Occupational Therapists, Physiotherapists and Orthotists with specialised training use several processes to manufacture custom made splints designed to meet the clinical requirement of a specific condition. Failure to produce a splint that meets these requirements, can impact recovery and can result in long standing deficits to the patient.

Failure to produce a splint that meets these requirements, can impact recovery and can result in long standing deficits to the patient. Ultimately, reduced function of the upper limb may impact on the patient’s ability to complete Activities of Daily Living (ADLs), maintain employment, purse leisure activities and even their participation in the community.The current clinical practice is to immobilise the target anatomy using a splint device. Immobilisation is typically achieved using plaster of Paris, which can be highly restrictive and uncomfortable for the patients. Splints can comprise either an “off the shelf” pre-made variant, or a custom device made from a thermoplastic material. The choice of which is implemented is typically873based on the judgement of the healthcare professional.

The choice of which is implemented is typically873based on the judgement of the healthcare professional. Fabrication of a custom made upper limb splint can be a complex process by the necessity to tailor requirements to the individuals needs with respects to anatomical variances, aesthetic appearance and cost. From the practitioner’s perspective, additional factors relating to cost, easy of manufacture and the suitability of available materials to achieve the desired end result play a part in the decision making process. Unfortunately, even with the provision of custom made splints, using the current techniques and materials the rate of non- adherence is relatively high, with comfort, splint appearance, influence on ADL’s being cited as areas of concern [3]. As the devices are also handmade, there is the issue of variability in the device and its effectiveness between different clinics, which further confound this issue.

As the devices are also handmade, there is the issue of variability in the device and its effectiveness between different clinics, which further confound this issue. Therefore the current accepted process for manufacturing upper limb custom made splints has the opportunity to be challenged, with new technology, design and processes implements to better meet the needs of the people who require these types of splints.Additive Manufacturing (AM) is an emerging fabrication methodology which allows for the construction of products using a systematic ‘additive’ layer by layer build process and in complex configurations that would not be otherwise attainable using traditional manufacturing processes [4, 5]. Indeed AM allows for the relatively straight forward realisation of highly complex digitally generated models, with minimal user inputs to the manufacturing process, or intimate knowledge of the materials they are fabricated from.

Such versatility in the build process is resulting in a disruption of traditional manufacturing paradigms, with impact in several major industries ranging across the automotive [5, 6] to the medical sectors [7-9]. With respect to the medical industry, AM’s potency to disrupt existing practises stems for the ability to reproduce even the most complex of human forms in what has been described as patient specific devices, with practical examples being demonstrated for patient centric bone replacement implants [8], prosthetics [10] and rehabilitation/splint based devices [11-14]. The potential of this technology is enhanced further by the alignment of technologies to rapidly digitally reproduce the complexity of the human form, with demonstrations using both optical surface scanning [15, 16] and medical imaging [7, 10] data sets.More recently, AM technologies have been utilised toward the development and production of custom splint devices [12, 15, 17-19].

Researchers are aiming to leverage the advantages found in non-contact data digitisation, advanced Computer Aided Design (CAD) and AM to realise more patient centric designs which overcome limitations of current devices and provide a less intrusive experience for the patient during the orthosis creation. With respect to current limitations of current splint devices, insights have revealed a range of factors relating to the non-adherence of wearing an orthoses [3, 18]. Factors include difficulties keeping splints clean/dry, poor aesthetical qualities, discomfort due to poorly fitting devices, compromised ability to perform routine tasks and odour related issues.

Factors include difficulties keeping splints clean/dry, poor aesthetical qualities, discomfort due to poorly fitting devices, compromised ability to perform routine tasks and odour related issues. It is believed that the versatility of CAD and AM can be applied to address these issues, realising more ergonomic and aesthetically appealing designs, tailored to the individual’s unique needs and priorities.In this study, we examine the use of high resolution optical surface scanning, CAD and AM to develop a new patient centric thumb splint device. We initially use optical scanning to rapidly capture a person’s hand anatomy while they maintain the position for the final form of the thumb splint. We then convert the captured data into a three dimensional digital model, which is used as the template to form the splint device.

We then convert the captured data into a three dimensional digital model, which is used as the template to form the splint device. We examine the ideal tolerances of the device alongside design features to ensure ease of fit, while still maintaining potential therapeutical effectiveness desired by a standard thumb orthosis. An open porous design is also examined to create an aesthetically appealing device, which more readily allows for the release of trapped moisture on the skins as compared to a tradition splint [12, 19-21]. Once a final design was derived, the model was manufactured using Fused Filament Fabrication (FFF) in ABS polymer to assess the final device and its potential for use in place of a874traditional thumb splint.

Overall, we believe we have developed a robust methodology for thumb orthosis design and manufacture in a form which is more ergonomic, provides better fit and aesthetic qualities.2 Existing Thumb Splint FabricationThe short thumb spica splint is a commonly used splint that can be used across a number of musculoskeletal conditions including acute ligament repair, bone fracture and osteoarthritis of the thumb. This splint was specifically chosen to challenge the technology being studied as the desired immobilisation position the hand is one of the more challenging orientations to achieve in a respective device. The thumb position for the splint is vital to ensure the correct anatomical position is achieved, whilst also allowing functional use of the hand by maintaining freedom of the metacarpophalangeal joints of the fingers and interphalangeal joint of the thumb, whilst not restricting the natural action of the wrist.

A low temperature thermoplastic (Aquaplast, OPC Health, Australia) is used to manufacture this splint, a material that is both strong and flexible enough to meet the requirements of this type of design, and the typical material used by the orthotist assisting in this study. An overview of the fabrication workflow can be seen in Figure 1.Figure 1: A diagram of the typical workflow performed by an occupational therapist to design and fabricate a thumb splint875In this study we asked a local orthotist to design a thumb splint for the volunteer in this study, which would act as a comparison for the device we will develop in this study. The technique described is a relatively typical example of how a thermoplastic thumb splint is created.

The technique described is a relatively typical example of how a thermoplastic thumb splint is created. The initial stage of the manufacturing process requires the health professional to create a template of the hand, this is commonly done by tracing an outline of the hand on paper/card and marking significant anatomical landmarks, such as the fingers, thumb, etc. A preliminary splint design is then traced over this template, and if the health professional is happy with the design, this is transferred onto the thermoplastic. The thermoplastic is cut to size and then heated in a hydroculator (heat bath) to the specified temperature of the thermoplastic, in this study this was approximately 70oC. The health professional then places the patient’s upper limb into the required position ensuring that the thumb, fingers and wrist are in the correct orientation.

The health professional then places the patient’s upper limb into the required position ensuring that the thumb, fingers and wrist are in the correct orientation. This position is often held by the patient for several minutes while the health professional removes the thermoplastic from the water, lets it cool to a safe temperature and then places this directly onto the patient to mould the splint to the patients hand. This process relies on the skill and knowledge of the health professional to be able to mould the splint as required, in addition to the ability of the patient to maintain the desired position. Once the thermoplastic has cooled, it begins to harden and can be removed from the patient for any additional alterations. Alterations can include cutting away of excess material, addition of straps to secure the splint or any other modifications as required.

Alterations can include cutting away of excess material, addition of straps to secure the splint or any other modifications as required. If further alterations are required, the thermoplastic can be reheated but there is a risk it may lose its shape.3 MethodologyThe primary elements of this study comprise the acquisition of a person data using optical scanning, construction of a 3D anatomical model of the hand, the development of the orthosis from this scan data, the manufacturing of the device and the final evaluation. A summary of a typical workflow can be seen in Figure 2.2.1 Optical ScanningSurface topography scans were obtained using a light reflectance scanning system (Spider, Artec, Luxembourg), which is minimally invasive and utilises visible light, meaning scans can be performed with no health risks to the person being scanned.

The scanner has a scan resolution of approximately 50-100µm, which allows for both the major anatomical form and minor details such as creases in the skins surface to be resolved. In this study the scanner was set to a resolution of 500µm which was considered suitable for the level of detailed required to be captured. The scanner is designed to work alongside a propriety software package (Artec Studio 10, Artec, Luxembourg) which streamlines the data acquisition phase and allows for real-time visualisation of the capture process.To capture a surface of an object, the scanner is activated and held as a constant distance of approximately 50cm from the object, before being translated around the object to build up a surface profile. Care must be taken to not translate the scanner too quickly (>10cm/s) or to move to far away from the optimal imaging distance or the scanner loses tracking of the object. A typical scan procedure can be completed within 3-5 minutes.

A typical scan procedure can be completed within 3-5 minutes. The scanner software also allows for the data processing of the scanned surfaces to remove spurious data points, noise and unwanted scan data. The software is also capable of rudimentary smoothing and fixing processes to create an enclosed 3D model, but generally does so with some degree of data loss of finer details. Therefore, such operations were performed independently using CAD software to retain the model details.876Figure 2: A diagram illustrating the workflow from data capture to final production of the thumb splint device.2.2 Computer Aided DesignThe scanner data was post processed using the CAD software 3-Matic STL 10.0 (Materialise, Belgium). Initially, the model was checked for digital errors before, some minor smoothing operations to improve the consistency of the data.

Initially, the model was checked for digital errors before, some minor smoothing operations to improve the consistency of the data. Following this the model was remeshed to improve the consistency of the overall mesh, before the data was suitable for construction of the thumb splint. The splint is realised using an approach of applying multiple surface projections to create hand models that were 0.5mm and 3.5mm thicker than the original hand model. The hollow form of the model was created using Boolean operations, followed by selective trimming to create the basic form of the model. In this study we examined an approach comprising a single piece splint design, which can be strapped to the person using a single piece of Velcro.

In this study we examined an approach comprising a single piece splint design, which can be strapped to the person using a single piece of Velcro. This matches the typical configuration of traditional thumb splint device.Beyond formation of the general shape of the device, we examined a technique of projecting a two dimensional pattern on the surface of the model, before extruding a three dimensional shape into the segments of the pattern as we have described previously [12, 19]. This variant of design was realised using 3-Matic STL (v10, Materialise, Belgium).8772.3 Additive ManufacturingOnce a device design had been finalised, the resulting models were 3D printed to create the final thumb splint model. Designs were manufactured using Fused Filament Fabrication (FFF) using a Flash Forge Creator Pro (Flashforge Corporation, Zhejiang, China) in standard ABS material, which provides adequate material stiffness to manufacture functional models.

When printing a hotend temperature of 232oC, bed temperature of 105oC and layer resolution of 0.2mm were employed. When printing, parts were orientated upright in the z-direction (Figure 2-3) to ensure the parts were printed with minimal support material and achieved the best cosmetic finish with minimal surface roughness.3 Results3.1 Patient specific data acquisitionThe primary intension of this study was to develop a methodology for streamlined data capture and model development and so scanning procedures were performed on a healthy individual who suffered no complications with respect to the thumb and its mobility. For therapeutic effectiveness, the correct orientation for immobilisation of the thumb is in a position which was described by the orthotist as the shape the hand would make ‘holding a soft drink can’.

A picture of the hand position can be seen in Figure 2-1, where the fingers are curved inwards in an arc shape, while the thumb holds a position parallel to the fingers, equally with an arc shape. The volunteer was asked to hold this position with their arm partially extended to allow for easy access of the scanner. Once in position the scanner was translated through various orientations, at a fixed distance, to construct the various surface profiles of the volunteer’s hand. It was found that the majority of the hand could be easily resolved with a single translation of the scanner. However, additional adjustment of the scanner capture acquisition time and image saturation had to be suitably adjusted to ensure reflection based aberrations were reduced to produce a well-defined image for capture.

It was also found that the inner contours of the hand proved to be more challenging to scan due to scan acquisition issues and the awkward positions that the scanner had to be rotated into to capture the lower portion of the hand. In this instance we believe scanning issues stemmed from interference of the reflected light due to natural shadows of the hand or reflectance of the light away from the scanner detectors. To compensate for these issues, several adjustments were made to the scanner to perform scan using the ‘fast fusion’ function, which disregarded skin pigmentation data, as only topological data was required.Following optimal scanner set up, the data acquisition phase was not without its own challenges. Firstly, minor movements of the volunteer’s hand resulted in alignment discrepancies during the post scan image processing. For the most part this was a minor issue as the scanner software was found to adequately compensate for this.

For the most part this was a minor issue as the scanner software was found to adequately compensate for this. Secondly, and perhaps more importantly, there was a certain degree of discomfort described by the volunteer in having to keep their hand and arm still for the duration of the scanning while we were optimising the best technique for data acquisition. Once our scanning techniques had been optimised, a typical scan could be performed within the space of 2-3 minutes. We therefore acknowledge that should this technique be implemented in real circumstances there will no doubt be a period of user training to refine optimal scanning technique by a practitioner. Comparing this to traditional methods of capturing a person’s anatomical data using thermal polymer moulding, the examined technique is both non-invasive, quicker and potentially provides greater accuracy in reproduction. We hope to further qualify these878differences in further studies.

We hope to further qualify these878differences in further studies. Ultimately, despite both techniques suffering some limitations, we concluded that the scanning approach was the superior of the two techniques.Figure 3a) illustrates segments of the scan data captured of the volunteer’s hand, where it was found that the final model could be adequately constructed from three separate scan data segments. It was also found that due to complications with data acquisition, the final form could not be adequately resolved to the point where an enclosed 3D model could be constructed, and so final finishing of the surface data was performed manually using CAD software.3.2 Orthosis DesignInitially the surface data of the volunteer was manipulated to completed regions that were unable to be captured during the scanning phase.

This was achieved in the 3-Matic software using a manual triangulation bridging, where regions of the model mesh were manually reconstructed to complete the form of the model. Figure 3b) shows the compiled raw surface data which was output from the Artec software alongside the final, trimmed 3D model of the hand.Figure 3: a) A picture of the scanner and image data processing equipment illustrating segments of the hand surface being stitched together and b) stages of the raw data from the scanner software to the final closed 3D model template data used for splint construction.879As a primary objective in this study was to create a more patient oriented splint, we use the model as the template to form the device. When manufacturing the splint, it is desirable to obtain a design which has adequate stiffness to resist the typical forces the patient may impose on the device to avoid breakages.

When manufacturing the splint, it is desirable to obtain a design which has adequate stiffness to resist the typical forces the patient may impose on the device to avoid breakages. Feasibly, the part could be deliberately made to larger thicknesses to increase the overall stiffness, thereby avoiding breakages. However, it is desirable to achieve a compromise such that the part is make as light and as thin as possible to increase the users comfort when wearing the device. From preliminary in house testing a thickness of 3mm was considered the ideal minimum thickness and so was used as the thickness tolerance for the final splint design.

From preliminary in house testing a thickness of 3mm was considered the ideal minimum thickness and so was used as the thickness tolerance for the final splint design. We hope to provide a more comprehensive evaluations of the stiffness characteristics resulting from varying part thicknesses in future studies.Figure 4: Various designs phases illustrating a) the thumb splint concept template and b) the final design of the projected pattern style splints.The initial stages of forming the hand splint comprised the trimming down of the hand model to form a template of the how the splint will be located on the hand. To guide the trimming process, the distal palmar and interphalangeal creases of the hand were used, where a clearance of 10mm from the distal palmar and 5mm from the interphalangeal creases were used to ensure that the potential device would not restrict movement of the upper portion of the thumb, fingers and bending about the distal palmar.

The lower portion of the splint was trimmed approximately 10mm above the crease of the wrist to ensure that when the splint was worn mobility was not compromised. From this template data, the surface of the models was extruded to 0.5mm and a second model to 3.5mm,880before a Boolean subtraction was performed to create a hollowed form of the splint. Then a region (65-70mm long and 25-35mm wide) relative to the back of the hand was trimmed and two recess holes created for the placement of a velcro strap. An image of the final template concept can be seen in Figure 4a).

An image of the final template concept can be seen in Figure 4a). The digital template was now ready for application of the porous pattern and additional design stylisations.Figure 5: Images of the manufactured splint designs a) post print and post processed, b) being worm and used by the volunteer and c) compared to a traditionally designed splint.881To construct porous design, an approach using a 2D projected pattern was apply, similar to what has been described previously [12]. Firstly, the model is segmented to leave a border of 5mm around the edges of the model, alongside a patch around the back of the thumb. This border ensured that the porous pattern would not be applied around the edges of the model, which could result in stress concentration points, in addition to points which could potentially press into the wearer’s skin.

The region behind the thumb was to act as a platform onto which we would incorporate an example of a custom design feature to demonstrate the mass customisation potential of such splints. In this instance we incorporated the Deakin University logo onto the thumb model, as can be seen in figure 4b). For the remaining portion of the splint, a 2D pattern was projected onto the surface of the template and adjusted to given a complex, graded size to the pattern, further exemplifying the versatility of this approach. Once the projection was achieved a generic circle pattern was then extruded throughout this pattern resulting in the final design found in Figure 4b).3.3 ManufacturingThe splint design was printed in ABS plastic using 100% infill to ensure the most robust part possible, and printed using a raft material to optimise print bed adhesion. As the part is to be worn, keeping the manufacturing surface roughness as low as possible was desirable to maximise comfort for the wearer.

As the part is to be worn, keeping the manufacturing surface roughness as low as possible was desirable to maximise comfort for the wearer. It was recognised that due to the complex form of the orthosis that support material would be required to ensure the integrity of the model during the build process. Despite this, initial tests were performed to examine if the design could be printed without the need for support material to both reduce the need for additional post process and to reduce additional material that could compromise the surface finish of the part, as the printer utilised the print material as the support material. All such printing attempts failed and so it was deemed necessary to use supports.It was unknown what configuration of support material would be ideal and so two commonly used configurations were tested comprising ‘rectangular’ and ‘tree like’ branching types.

Our settings for printing supports comprised overhang angles less than 40o and with a structure thickness of 3mm. Both configurations provided adequate support to reproduce the part in its entirety. Of the two structures, the tree like supports minimised the contact area with the part and were found to the easiest to remove and so were considered the preferred choice of supports. It was found that in some instances, when the supports were removed there was residual material left behind on the model. To remedy this, a craft knife was used to scrape the bulk of the material off, followed by a light sanding using 120 grit sand paper, until a smooth finish was achieved. Following this methodology, a part was ready for use within 3-5 minutes following printing. Figure 5 illustrated the 3D printed part as it appeared of the printer, following post processing and in its final form containing the Velcro strap and placed upon the wearer’s hand.

Figure 5 illustrated the 3D printed part as it appeared of the printer, following post processing and in its final form containing the Velcro strap and placed upon the wearer’s hand. As a comparison, we fabricated a thermoplastic splint using traditional methodologies, which can be seen next to the developed 3D printed splint in figure 5c). Superficially, the shapes of both matched one another, resulting in the hand of the wearer being orientated in approximately the same position. From a qualitative external inspection it can be seen that the 3D printed splint only uses the material necessary for production. By contrast the traditional splint contains sizable amounts of excess material folded over the outer surfaces around the opening for the thumb and the area in contact with the palm of the hand.

By contrast the traditional splint contains sizable amounts of excess material folded over the outer surfaces around the opening for the thumb and the area in contact with the palm of the hand. Additional design benefits comparing the devices in this study are that the 3D printed design incorporates openings for the velcro strap, where by contrast the traditional splint required adhesive Velcro pads to incorporate the strap. Indeed the velcro fastening mechanism highlights additional limitations to the existing design. Firstly, the adhesive pads for the Velcro are often prone to loosing adhesion when exposed to repeatedly to moisture, causing failure of the splint. Additionally, the velcro strap only has two small contact points882to maintain the tension of the device. By contrast the 3D printed splint, with the velcro strapping round the persons hand does not require any adhesive pads and can accommodate a much larger region of velcro hooks to keen the strap in place.

Ultimately this leads to a much more firm fit which is less prone to failure. Further comparisons of the two designs can be found in table 1. It can be seen that both splints have identical weights but the new splint offers greater stiffness and a 3-12 factor increase in porosity, meaning a greater capacity for moisture release. In principal, the higher the porosity, the greater the capacity for a splint to release moisture, neglecting a materials moisture wicking capability. We hope in future work to develop concepts to investigate elements of moisture release in greater detail. It is worth noting that a negative to the 3D printing process is the fact that the splint took approximately eight to nine times longer to manufacture, which is still acceptable but less than desirable with respect to it potentially meaning an additional visitation by a potential patient.

We believe that as printing technology develops, that the turnaround time for a part will begin to decrease and so this limitation may be overcome in the near future for FFF printing technology. Alternatively, new forms of printing could be employed, such as Continuous Liquid Interface Production (CLIP), which possess greater print resolution, with vastly improved speed of printing, potentially reducing production time by a factor of up to 50.DesignTraditional Splint Hole Patterned SplintManufacturing time (hrs) 1 8-9Weight (g) 30 30Approximate Porosity (%) 2-8 24Table 1: Comparison of the various thumb splint manufacturing times and properties taking into account data capture, design, manufacturing and post processing times.3.4 Qualitative AssessmentIn a recent review conducted by Kelly et al 2015, various user based metrics were described relating to the adherence of wearing an upper body splint.

Factors included difficulties in maintaining cleanliness and hygiene, aesthetics, comfort and ergonomics. However, such factors are arguably qualitative in nature and relate to the individual users experience and perception of the device. Therefore, performance metrics relating to such devices can only be qualitative in nature. In an attempt to assign metrics to assess such factors with respect to splints, research groups have taken an approach of user-centred experience questionnaires to either validate designs or provide suggestions for improvements [22]. In this study we apply a similar methodology to assess the effectiveness and ergonomics of the thumb splint. We asked the volunteer of the study to wear each device for a period of 2-3 minutes and perform tasks to bend the palm of the hand, attempt to move their thumb in a circular motion, bend their thumb, rotate the wrist and finally to perform a pincer action with the thumb and fingers, as can be seen in figure 5b).

Once motions had been complete we asked the volunteer to rank each device on a scale of 1 to 10 (1 being low and 10 being high) for a series of user centred questions. These included:Q1) How do you rate the aesthetic qualities of the splint?Q2) Would you feel happy wearing this device in public?Q3) How comfortable is the device to wear?883Q4) How stiff would you say the device feels when worn?Q5) How much mobility of your hand would you say you have remaining to perform everyday tasks?Q6) How would you rate the device for ease of cleaning your hand and for release of moisture while being worn?Following the questions we would obtain a score for each device with a maximum of 60. Based on these findings the highest scoring device would be considered optimal. Results for each question are found in Table 2, where it can be seen the 3D printed splint scored highest, and is therefore considered optimal in this study.

Results for each question are found in Table 2, where it can be seen the 3D printed splint scored highest, and is therefore considered optimal in this study. It is noted that the results are highly subjective and only based on the single user analysis and so findings will no doubt vary between larger cohorts of users. We hope to develop these initial questions, suggested hand motion evaluations of the participant, wear time and user metrics, and to apply them to a larger cohort size in future studies. However, these very preliminary findings hint at the potential of AM devices to offer user evaluated improvements over traditional devices.Design Traditional Splint Projected Patterned SplintQ1 3 7Q2 5 9Q3 6 8Q4 4 8Q5 6 8Q5 2 6Total Score 26 46Table 02: User based qualitative assessment of the various splints4 ConclusionFindings from this study have proven the efficacy of using non-invasive optical scanning to readily derive patient specific anatomy.

Following a brief period to refine our scanning techniques, the digitised data can easily be acquired using the Artec system, converted into a 3D hand model and be used as template from which to form a patient specific thumb splint concept. We have found there to be considerable scope for design freedom, and we demonstrate the efficacy of creating a design incorporating a generic, porous pattern. The design was evaluated by a trial user and stated to provide an excellent fit, while being both an aesthetically pleasing and ergonomic configuration to allow for ease of hand mobility. We note that the user evaluation is largely opinion based and so to more accurately affirm the efficacy of the splint a significantly larger cohort study (>25 people) must be performed, and done so over extended periods of user wear time. It is also noted that the turnaround time for fabrication of the new device is several hours longer than for the traditional device.

It is also noted that the turnaround time for fabrication of the new device is several hours longer than for the traditional device. We believe that this challenge will be addressed over time as printing technology develops. However, more importantly the improvements AM offers may lead to greater user adherences of such devices which ultimately would outweigh the negatives of longer fabrication time with regards to patient outcomes. We hope to better understand this facet in future studies and encourage other groups to explore the potential of AM for upper limb splints towards improved patient adherence.AcknowledgementsThis research was conducted by the Australian Research Council Industrial Transformation Training Centre in Additive Biomanufacturing (IC160100026). All resources for the study were884provided courtesy of the School of Engineering at Deakin University.

All resources for the study were884provided courtesy of the School of Engineering at Deakin University. Ethics approval for this project was granted by the Deakin University Ethics Committee (HEAG-H 55_2018) following conformity to the requirements of the National Statement on Ethical Conduct in Human Research (2007).References[1]M. Ahern, J. Skyllas, A. Wajon, and J. Hush, "The effectiveness of physical therapies for patients with base of thumb osteoarthritis: Systematic review and meta-analysis," Musculoskeletal Science and Practice, vol. 35, pp. 46-54, 2018/06/01/ 2018. H. J. Benjamin and B. T. Hang, "Common Acute Upper Extremity Injuries In Sports," Clinical Pediatric Emergency Medicine, vol. 8, pp. 15-30, 2007/03/01/ 2007. L. O'Brien, "Adherence to therapeutic splint wear in adults with acute upper limb injuries: a systematic review," Hand Therapy, vol. 15, pp. 3-12, 2010/03/01 2010. M. I. Mohammed and I.

Collins, and I. Gibson, "Design Optimisation of a Thermoplastic Splint," in Proceedings of the 28th Solid Freeform Fabrication Symposium, Austin, TX, 2017, pp. 2409-2418. A. Paterson, R. Bibb, and R. Campbell, "Orthotic rehabilitation applications," Medical modelling: the application of advanced design and rapid prototyping techniques in medicine. New York: Woodhead Publishing, p. 283, 2014. A. A. Portnova, G. Mukherjee, K. M. Peters, A. Yamane, and K. M. Steele, "Design of a 3D- printed, open-source wrist-driven orthosis for individuals with spinal cord injury," PLOS ONE, vol. 13, p. e0193106, 2018. P. Volonghi, G. Baronio, and A. Signoroni, "3D scanning and geometry processing techniques for customised hand orthotics: an experimental assessment," Virtual and Physical Prototyping, vol. 13, pp. 105-116, 2018/04/03 2018.885[16] M. J. Pilley, C. Hitchens, G. Rose, S. Alexander, and D. I.

viewed PaperDesign and Additive Manufacturing of Bio-Inspired Copper Heat Sinks for Microelectronics CoolingJordan Yaple1, Cameron Noe1, Abdulmajeed Alenezi 2, Patrick Phelan2, Dhruv Bhate113DX Research Group, Arizona State University, Mesa, AZ 85121 2School of Matter, Energy and Transport Engineering, Arizona State University, Tempe, AZAbstractThe heat sink is a key component in thermal management of microelectronics and is traditionally designed with arrays of pins or fins. In this work, a bio-inspired approach to designing heat sinks was evaluated. A process was first developed on a 100W Laser Powder Bed Fusion (LPBF) system for the copper alloy CuNi2SiCr, which was found to have a tenfold deficit in thermal conductivity relative to commercially pure Copper.

This process was then used to fabricate four heat sinks: two replicates of commercial heat sink designs, and two leveraging bio-inspired design approaches with the aim of increasing the available surface area per unit volume. The four designs were tested on a microelectronics package placed on a hot plate and their performance compared against commercially available heat sinks. The results demonstrated that despite their poor thermal conductivity, the bio-inspired heat sinks had equivalent performance to the commercial heat sinks.IntroductionOf all the alloys developed for the Laser Powder Bed Fusion (LPBF) process, copper alloys have emerged only recently as a viable metal to fabricate parts from. LPBF of copper has had several challenges associated with it, stemming primarily from its low absorbance in comparison of incident laser energy relative to other well-qualified LPBF alloys [1].

LPBF of copper has had several challenges associated with it, stemming primarily from its low absorbance in comparison of incident laser energy relative to other well-qualified LPBF alloys [1]. In addition to processability challenges, risk of damage to optical components [2], oxidization of copper [3], relatively low density [4] have also been identified as potential issues with the LPBF of copper. Once a sufficiently dense part is created, there are limitations in an application context from excessive surface roughness [3], [5]. Nonetheless, prior work has demonstrated the feasibility of fabricating components using LPBF of copper alloy [6], and this forms the basis for the work here, which is an exploration of the use of LPBF and a specific copper alloy (CuNi2SiCr) for the fabrication of heat sinks for microelectronics cooling, previously published as a Master’s thesis [7]. The three main objectives for this work were:1.

The three main objectives for this work were:1. Development of a working process for LPBF for CuNi2SiCr with an as-printed density at least exceeding 94%. This criterion was based on prior literature where densities in excess of 94% were shown to be feasible.2. Manufacture and compare the cooling effect of commercially available (COM) heat sinks against nominally identical additively manufactured (AM) heat sinks.3. Design, fabrication, and evaluation of bio-inspired heat sinks.1826The flow of this paper is as follows: the designs are first discussed, including the motivation for the bio-inspired designs, followed by the manufacturing aspects of this work, including the process development for the copper alloy, followed by the fabrication of the heat sinks. Testing and characterization results are then reported, followed by a discussion of these results.

Testing and characterization results are then reported, followed by a discussion of these results. The paper ends with a conclusion summarizing the key findings of this work.Heat Sink DesignTwo commercially available heat sinks were first acquired: one pin-based and one fin- based. The heat sinks that were chosen to fit on top of a microelectronics heat spreader on a 37.5 x 37.5mm substrate typical for CPUs. A combination of design drawing data and actual measurements with calipers were used to model these components in SolidWorksTM [8] and used to create a baseline comparison for the cooling effect between additive and commercial parts for the same nominal design. These designs are shown in Figure 1a and 1b for the commercial pin- and fin-based heat sinks, referred to henceforth as COM Pin and COM Fin.Figure 1.

These designs are shown in Figure 1a and 1b for the commercial pin- and fin-based heat sinks, referred to henceforth as COM Pin and COM Fin.Figure 1. The four different heat sink designs explored in this work: (a) pin, (b) fin, (c) Cholla cactus inspired, (d) Schwarz-P TPMS, and (e) the cholla cactus skeleton that served as inspiration for one of these designs1827The second pair of designs followed a bio-inspired approach. The first of these, shown in Figure 1c, was inspired by the skeleton of the cholla cactus, shown in Figure 1e. This particular organism was chosen for its lightweight structural performance [9], and also its ability to survive in extreme climates.

This particular organism was chosen for its lightweight structural performance [9], and also its ability to survive in extreme climates. In particular, the perforations along the cylindrical structure suggests an ability to enhance convection by enabling access to the interior surface and potentially improving overall heat sink performance, which has been previously shown to improve airflow for both natural and forced convection for similarly shaped structures [10]. For this work, perforations were placed along a spiral path, following the structure of the cholla cactus, also designed in SolidWorksTM. The second bio-inspired design leveraged Triply Periodic Minimal Surfaces (TPMS) design which have now been well studied for a range of applications, and while they owe their origins to mathematical formulation, have also been found in several biological structures [11].

TPMS structures are essentially cellular materials, and span a wide range of application domains, and design exploration of this domain has grown significantly with the advancements in additive manufacturing [12], and a bio-inspired design approach is often employed when selecting a particular cellular material for a specific application [13] For this work, the Schwarz-P TPMS structure was selected, based on its high surface area to volume ratio, and the efficient thermal performance of these structures as reported in the literature [14]. These structures were designed in the Platform software developed by nTopologyTM [15].ManufacturingThe 100W Concept Laser MLab LPBF system was used in this work. This machine is considerably lower in maximum power than those typically used in the literature [1], [16], [17], so the first step was to establish a process that could perform at under 100W of laser power.

Using the optimal parameters found in other papers working with the CuNi2SiCr alloy, an initial process window was defined for scan speed, hatch spacing, laser power to achieve similar Volumetric Energy Densities (VED) as those reported in the literature. This was used to narrow a set of parameters for further refinement, summarized in Table 1. A total of 16 design points spanning the range of parameters shown in Table 1 were evaluated by building 10mm cube specimens on a 4x4 grid as shown in Figure 2a and 2b. The specific levels for each parameter varied across the range based on preliminary findings not reported here, in order to obtain a wide spread of VED. The cross section of each cube also shows the as printed porosity as seen in Figure 2c, under a microscope. A concern in the development of the process was the risk for oxidation of the copper powder, and the likelihood that this could result in degradation of printed density over time.

A concern in the development of the process was the risk for oxidation of the copper powder, and the likelihood that this could result in degradation of printed density over time. As a result, the same set of 16 cubes with identical process parameters were fabricated over a four-week period, beginning with virgin powder that immediately went from an unopened container into the MLab dose chamber and was used for fabrication within 4 hours, during which time the machine was inerted with Argon gas. Once the print was completed, the powder was left in the machine for a week at a time (in ambient conditions), and the same build repeated. The 64 specimens obtained were characterized using the Archimedes density method [18] within two days of being printed (including the wire-EDM process to remove the specimens from the build plate), and the results obtained are shown in Figure 3.1828Table 1.

Archimedes density for the 16 cubes as a function of VED and the exposure time, showing no significant changes with time at least for the first 2 weeks of exposure with the red lines indicating the average density.The first thing of note in Figure 3 is that the density range for the specimens tested spanned approximately 7.8 to 8.5 g/cc, whereas the reported density for this alloy from one supplier’s datasheet is 8.85 g/cc [19], indicating that the process developed with the 100W laser has room for improvement. The second thing of note is that the density did not show significant degradation, at least over the first 14 days, with slight reduction observed after 21 days.

The second thing of note is that the density did not show significant degradation, at least over the first 14 days, with slight reduction observed after 21 days. This data therefore suggests that oxidation risks for CuNi2SiCr, at least in the context of density, are manageable and for this particular study set a 2-week limit on fabricating parts with freshly opened powder.Based on these findings, an optimum process parameter set was selected as the one that gave the highest average density across all four weeks, which amounted to 94.3% relative density, with the datasheet value [19] used as a baseline. It is important to note that this reported value was not from the same equipment or powder supplier. Nonetheless, this was the highest relative density value attainable within the period of study on the Concept Laser MLab 100W, and a cross-section of one of the samples that employed these parameters is shown in Figure 3c.

The relatively slow scan velocity is indicative of the need for high VED for this process.Layer thickness: 0.03mm • Laser power: 85 W • Scan velocity: 47 mm/s • Hatch spacing: 0.1 mm • Continuous scan patternIn addition to Archimedes density method, two cylinders (25mm tall and 11mm diameter) were fabricated for estimation of thermal conductivity using the hot-disk sensor method [20]. Their bases were polished prior to testing. The resulting thermal conductivity measurement for the LPBF-fabricated CuNi2SiCr specimens was a mere 30 W/mK, substantially less than the 350-390 W/mK reported for commercially pure Copper, and a third as much than that reported by the1830datasheet obtained, where the as-printed thermal conductivity was stated to be 90 W/mK, which increased to 190 W/mK after precipitation hardening [19].Once these parameters were finalized, heat sinks were fabricated with them and are shown collectively in Figures 4a-4f.

The AM Fin (Figure 4a) and AM Pin (Figure 4c) heat sinks were manufactured using the same design as the two COM heat sinks (Figures 4b and 4d) for a comparison to isolate the difference in thermal performance between pure copper commercially fabricated heat sinks and the CuNi2SiCr sinks fabricated with LPBF. Following these prints, the Cactus inspired, and Schwarz-P heat sinks were fabricated with LPBF and are shown in Figure 4. Similar to the AM parts, the bio-inspired heat sinks both had 5mm of copper added to the base for base plate removal tolerance. There was a 5mm allowance added to the bottom of the heat sinks so they could be removed from the build plate with a wireEDM without cutting too deep into the heat sink base.

There was a 5mm allowance added to the bottom of the heat sinks so they could be removed from the build plate with a wireEDM without cutting too deep into the heat sink base. The specimens were then ground down to match the cross-section thickness of the COM heat sinks prior to testing.Heat Sink Characterization and TestingThe first observation of the differences between the COM and AM heat sinks is the surface roughness, with AM heat sinks being significantly rougher, as expected for LPBF, as shown in Figures 5a-5f. To quantify this roughness, regions were selected within the four AM and two COM heat sinks, as shown in the green boxes in Figures 5a-5f, and scanned under the Keyence VR-3200 structured white light scanning microscope [21] using the in-built curved surface correction algorithm to obtain an Sa value. The obtained roughness values from the six heat sinks are given in table 2.

The obtained roughness values from the six heat sinks are given in table 2. While the AM heat sink roughness values are highly dependent on orientation of the scanned region relative to the build platform, they nonetheless are indicative of the differences between commercially fabricated, and as-printed LPBF heat sinks with regard to surface morphology – this difference is very significant for the pin-based heat sinks, but not as significant for the fin-based sinks, since the commercial fin-based sink had a far more striated texture (Figure 5d) compared to the smooth pin texture (Figure 5b). The AM heat sinks were also measured using calipers and all measurements were found to be within 10% of the nominal design, for the commercial and bio-inspired heat sink designs.1831Figure 4. Additive manufactured heat sinks next to commercially available heat sinks: (a) AM Pin, (b) COM Pin, (c) AM Fin, (d) COM Fin, (e) Schwarz-P, (f) cactus inspiredTable 2.

Additive manufactured heat sinks next to commercially available heat sinks: (a) AM Pin, (b) COM Pin, (c) AM Fin, (d) COM Fin, (e) Schwarz-P, (f) cactus inspiredTable 2. Surface roughness (Sa) value comparisons for the AM and COM heat sinks (all in µm)AM – PinCOM – PinAM – FinCOM – FinCactus5795947781832Schwarz-P90Figure 5. Magnified images of the surface finish of each heat sink and the measurement location used for surface roughness calculation: (a) AM Pin, (b) COM Pin, (c) AM Fin, (d) COM Fin, (e) Schwarz-P, (f) cactus inspiredTo estimate the cooling performance of the COM and AM heat sinks, a test setup was built as shown in Figures 6a and 6b, following the setup demonstrated by Constantin et al. [6]. Results obtained with forced convection were highly variable and as a result only natural convection results are reported here.

[6]. Results obtained with forced convection were highly variable and as a result only natural convection results are reported here. Multiple thermocouples were epoxied to the heat sink and CPU as well as a thermal camera used to detect any hot spot anomalies (none were found and hence not reported here). A hot plate was set to three temperatures for this study: 65, 100 and 150 degrees Celsius. A 37.5mm CPU package was attached to the hot plate using thermal tape. A thermocouple was attached to the surface of the hot plate using Kapton tape to measure the surface temperature of the hot plate during testing. A groove was ground into the CPU heat spreader (as shown in Figure 6c), where a thermocouple was epoxied, and the surface was then sanded smooth.

A groove was ground into the CPU heat spreader (as shown in Figure 6c), where a thermocouple was epoxied, and the surface was then sanded smooth. The cooling1833effect was defined as the temperature of the CPU thermocouple with and without a heat sink on top of it, after allowing for stabilization of the temperature reading, which took a maximum of 15 seconds (consistent with [10]), which was then set as the standard wait time before recording a measurement. The other thermocouple measurements were used for reference and comparison purposes but not reported here. These measurements were made at three different temperatures – three measurements were made at each condition to obtain an average.Figure 6.

These measurements were made at three different temperatures – three measurements were made at each condition to obtain an average.Figure 6. Test setup for measuring cooling effect on account of heat sink: (a) CAD model of the experimental setup, (b) actual test setup used, and (c) groove in heat spreader used to mount thermocouple to obtain measurements right under the heat sinkTo first assess the differences between AM and COM heat sinks for the same nominal design, the average change in temperature due to the addition of a heat sink relative to no heat sink is plotted in Figure 7a, at each of the three temperatures for the fin and pin heat sinks. The commercial heat sinks had better in comparison to the AM heat sinks, and the COM pin has the best performance (i.e. highest change in temperature). While neither AM heat sink can compete with the best performing COM heat sink, this is to be expected given the ten-fold reduction in thermal conductivity.

highest change in temperature). While neither AM heat sink can compete with the best performing COM heat sink, this is to be expected given the ten-fold reduction in thermal conductivity. However, the fact that both AM heat sinks out-perform the worse COM heat sink with lesser sensitivity to the presence of forced convection is suggestive of the fact that the higher surface roughness in AM specimens disrupts the flow of air, resulting in a greater change in temperature.To assess the differences in design, the COM Pin design was chosen as a representative baseline since it was the design basis for the bounding box envelope of the bio-inspired structures. The data comparing cooling effects for the bio-inspired heat sink designs against the COM and AM pin heat sinks is shown in Figure 7b.

The data comparing cooling effects for the bio-inspired heat sink designs against the COM and AM pin heat sinks is shown in Figure 7b. At higher temperatures in particular, it is evident that the bio-inspired designs significantly close the gap between the AM-pin and COM-pin – this is a remarkable outcome given the low thermal conductivity of the LPBF CuNi2SiCr used in this work.1834(a)(b)Figure 7. Change in temperature for (a) the four pin- and fin-based designs (b) the bio-inspired heat sinks in comparison to the COM and AM pin heat sinksDiscussionA key finding of this work is that despite their far lower thermal conductivity, the bio- inspired CuNi2SiCr heat sinks were close in performance to the commercial, pure Copper heat sinks under natural convection conditions. The reasons for this can therefore be attributed to the design and morphology of the bio-inspired heat sinks.

The reasons for this can therefore be attributed to the design and morphology of the bio-inspired heat sinks. An examination of the specific surface area (nominal surface area divided by the mass of the heat sink) for all six heat sinks is shown in Figure 8, clearly indicating the higher values for the bio-inspired designs, which is at least a portion of the explanation for the improved performance. A second contributor to this performance is potentially the higher surface roughness which may play a significant positive role by mitigating against the formation of an insulating boundary layer.Figure 8.

Specific surface area for all six heat sinks used in this work, clearly showing the bio-inspired designs having the highest values1835ConclusionsThe findings of this work can be summarized in each of the three objectives defined at theoutset:An LPBF process was developed on a 100W machine for CuNi2SiCr that resulted in 94% dense specimens with thermal conductivity a tenth of commercially pure copper • LPBF CuNi2SiCr heat sinks when compared to commercial heat sinks made of pure copper showed different performances: additively manufactured fin-based heat sinks outperformed the commercial counterpart, but the pin-based heat sink underperformed, with regard to cooling effect under natural convection conditionsFinally, despite the very low thermal conductivity, bio-inspired designed heat sinks had comparable performance to the commercial pin-based heat sink, attributable to their higher specific surface area and surface roughness – more work is needed to establish these relationshipsThis work adds to the body of work demonstrating the benefits of designing for AM to enhance performance.

SIGN AND ANALYSIS OF A TOPOLOGY OPTIMIZED TRANSTIBIAL PROSTHETIC SOCKET USING COMBINED STATIC GAIT ANALYSISR. M. Carbonell and R. H. CrawfordWalker Department of Mechanical Engineering, The University of Texas at Austin, TX 78712AbstractThis paper presents the design and analysis of an optimized transtibial prosthetic socket developed using the ground structure method of topology optimization (GSM). The socket wall between the distal 25% of the original socket and a proximal brim is replaced with an optimized truss geometry and a thin wall (1 mm). Separate trusses are developed for the loading conditions of three critical stances: heel strike, vertical (standing), and toe-off. The truss models are combined with critical components to create the final design. The proposed socket is 81.58% of the original socket volume and is designed for manufacturing using Selective Laser Sintering (SLS) and nylon- 12.

The proposed socket is 81.58% of the original socket volume and is designed for manufacturing using Selective Laser Sintering (SLS) and nylon- 12. The socket design is analyzed, with the material properties for sintered nylon-12, at 10% increments between heel strike and toe-off to determine the viability of both the socket and the corresponding methodology. Simulation results indicate that the design exceeds requirements for all tested stances.IntroductionTranstibial prosthetic limbs are used as a means of ambulation by an estimated 380,000 people in the United States alone (Stevens et al., 2019). Transtibial (below the knee) amputations have increased over the past thirty years primarily due to type two diabetes and motorcycle accidents (Sanders & Fatone, 2011). Regular use of existing, commercially available transtibial prostheses is complicated by users’ experience of pain, injury, and psychological distress.

Regular use of existing, commercially available transtibial prostheses is complicated by users’ experience of pain, injury, and psychological distress. Current research shows that the prosthetic socket, the interface between the residual limb and the prosthesis, is the component related to the majority of user complaints (Paternò et al., 2018).This research proposes a transtibial prosthetic socket developed using the ground structure method (GSM) of topology optimization. Critical components, both a brim at the proximal end and the distal 25% of the socket comprising a socket base, are integrated with a 1 mm thin wall and an optimized truss structure based on the numerical approximation from (Dorn et al., 1964) of the truss optimization methods originally described in (Michell, 1904). The socket is designed for fabrication with Selective Laser Sintering (SLS).

The socket is designed for fabrication with Selective Laser Sintering (SLS). This process is chosen due to its capabilities to fabricate long overhangs and thin walls found in this design, and prior vetting as a viable construction method for lower limb prosthetic sockets, specifically for sockets with adaptive geometric features built into their structure (Faustini, 2004; Pham et al., 1999; Rogers et al., 2001). This socket is not user ready, but instead serves as a proof-of-concept meant for bench testing. The purpose is to assess the viability of truss-based optimization for complex loading problems and determine its potential for modifying prosthetic socket topologies.

The purpose is to assess the viability of truss-based optimization for complex loading problems and determine its potential for modifying prosthetic socket topologies. Further work is needed to modify this design for clinical use.2063In a review of socket technologies, (Paternò et al., 2018) identified four primary socket failure modes: volumetric fluctuation of the residual limb, including both diurnal fluctuations and semi-permanent changes, which leads to discomfort and shifts in limb position relative to the socket; overheating and the resultant perspiration-related skin conditions; pressure or normal forces, which cause discomfort and injury; and shear stress between the limb and socket (or liner) causing damage to the skin (Paternò et al., 2018).Historically, little has changed about the topology of transtibial prosthetic sockets.

Since the first transtibial socket was developed by Pieter Verduyn in 1696 (Thurston, 2007), the primary changes have been either the material of the socket or the method of suspension. In the 20th century, three structural designs of socket with similar shell models emerged as the primary transtibial prosthetic sockets. These sockets vary in topology to adjust the locations of pressure on the residual limb (Bowker, 1992; Moo et al., 2009). The first is the total surface bearing (TSB) socket, fitted tightly to the body in an attempt to evenly distribute pressure on the residual limb. Next is the patellar tendon bearing (PTB) socket, developed in 1950 at the University of California at Berkeley. The PTB socket includes pressure tolerant regions in the area of the patellar tendon and the posterior of the socket. These are adjusted to bear more of the load within the socket, relieving pressure on the anterior of the limb and the fibular head.

These are adjusted to bear more of the load within the socket, relieving pressure on the anterior of the limb and the fibular head. It is common for a socket to be a hybrid of these two designs in an attempt to balance the benefits of each (Stevens et al., 2019). Least common of the three, but still widely accessible, are hydrostatic (HTS), or pressure cast, sockets that have consistent pressure on all parts of the residual limb but incorporate a pressure chamber to keep constant surface pressure as the limb adjusts due to soft tissue deformation. The HTS socket relies on the assumption that changing pressure points leads to even pressure absorption as the user changes positions; however, dynamics are not accounted for in the design.The increase in transtibial prosthesis users has led to an increased interest in socketimprovements both in academia and industry.

Ongoing, widespread user dissatisfaction with transtibial sockets bolsters continued research and calls for explorations of structural innovation in this field (Pirouzi et al., 2014). In literature proposing new topologies for transtibial sockets, additive manufacturing (AM), along with computer-aided design and manufacturing (CAD & CAM) are often critical to the realization of novel designs. SLS and Stereolithography are the most commonly used AM techniques in prosthetic socket development. Double walled sockets, solid sockets with varying wall thicknesses, and sockets with geometrically complex features are possible through AM and account for a large portion of proposed topology changes in transtibial socket research (Faustini, 2004; Faustini et al., 2006; Rogers et al., 2008; Webber & Davis, 2015).Each socket is customized due to variation in the shape of the residual limb. Prior to anymodification, the topologies are complex.

Prior to anymodification, the topologies are complex. Current methods of fabrication include a prosthetist casting or scanning the residual limb geometry and creating a positive mold for forming or laminating the socket materials. These molds are typically made of plaster and carved either by hand or using a specialty computer numerical control (CNC) lathe. Some molds are created additively from computed tomography (CT) scan data of residual limbs. Thus far, this has not been found to improve the fit of sockets, but it does reduce labor and allow molds to be created more quickly.

Thus far, this has not been found to improve the fit of sockets, but it does reduce labor and allow molds to be created more quickly. Customization, complex geometry, and high costs, due in part to labor, make this area of study an opportunity for AM sockets of practically any design to be created at2064comparable prices more easily (Hsu et al., 2018; Rogers et al., 2007; Sewell et al., 2012; Trower, 2006).In this research, there is particular interest in prosthetic socket applications that utilize AM to change socket topology or employ topology optimization in the design process. However, it should be noted that there are examples of innovative prosthetic socket structures, both in academia and industry, that are not fabricated using AM ("Martin Bionics," 2022; Quezada, 2017; Vaughan, 2014).SLS sockets with compliant features in regions of high pressure are described in (Faustini et al., 2006) and (Rogers et al., 2001).

The first version developed by (Rogers et al., 2001) used thin walls in high pressure regions. Building on this work, (Faustini et al. 2006) implemented an optimized spiral topology with bracing that served as a spring in these same regions of high pressure. This feature further decreased stiffness in these areas while still maintaining contact with the residual limb. A comparison with the thin wall design was conducted in (Rogers et al., 2008) with a single user. It was found that the user experienced increased comfort during ambulation and was able to increase their walking speed with the optimized pressure relief spring model (Faustini et al., 2006; Rogers et al., 2008; Rogers et al., 2001).Temperature change is the failure mode of transtibial prosthetic sockets that is addressed the least in current literature.

This issue can lead to sweating, skin maceration, and slipping in the socket, making temperature shifts highly uncomfortable and dangerous to the user (Ghoseiri & Safari, 2014; Paternò et al., 2018; Webber & Davis, 2015). An interesting approach to this issue is a design from (Webber & Davis, 2015). Using SLS, they created a socket with a passive helical cooling channel that wound around the residual limb. In bench testing, sensors were placed at the distal and proximal ends of the socket. At 31.1° C, the average delta from the outer wall to the inner wall was 4.5o C at the proximal sensor location and 5.1° C at the distal sensor location.Recently, examples of topology optimization and predictive design have entered the transtibial prosthesis literature. The first of these was a study described in (Steer et al., 2019, 2020).

The first of these was a study described in (Steer et al., 2019, 2020). This study used a multi objective genetic algorithm (MOGA) and surrogate limb models of varying geometry to determine areas appropriate for socket wall removal or reduction. The fitness functions used in this research sought to minimize distal loading and bony prominence loading (loading in areas with the least soft tissue between the residual skeletal structure and the socket). In these models, areas for socket wall removal or reduction were found for each limb geometry for biased optimization towards each fitness function as well as unweighted fitness functions (Steer et al., 2019, 2020).Another example of a computational design method can be found in (Ballit et al., 2022).This socket generation method was based on CT scan data and soft tissue deformation throughout ambulation. The socket design was constrained by force translation and the pain threshold for soft tissue.

The socket design was constrained by force translation and the pain threshold for soft tissue. Wall thickness was added or removed from areas of the socket model until an optimal socket was obtained. The final model was found to be below the soft tissue pain threshold throughout ambulation (Ballit et al., 2022).2065The final relevant example is the light-weighting of a socket base using solid isotropic microstructure (or material) with penalization (SIMP), described in (Sotola et al., 2020). The model of the socket base that was optimized attaches below the socket, providing additional support. However, while this component is preferred by some transtibial amputees, it alters the center of gravity in the leg in a way that causes discomfort and malignant gait adaptation that can pull on the leg, causing a pistoning motion that increases shear stress.

Keeping the exterior shell unmodified, the interior of the base was optimized through SIMP using multiple element shapes to compare results. The final design maintained all necessary features for integration with the transtibial socket along with the stiffness necessary to support the limb. Using different element shapes, a maximum reduction in weight of 58% was achieved for the most ambitious design that still met stiffness criteria (Sotola et al., 2020).These optimized sockets and socket components have yet to be manufactured as of thiswriting. However, it is likely these designs will utilize an additive process to vary wall thickness.

However, it is likely these designs will utilize an additive process to vary wall thickness. Due to the complex geometry generated through SIMP, and the desire to maintain the exterior shell, (Sotola et al., 2020) specifies plans to use SLS to fabricate their socket base in future work.MethodsThe methods for creating this proof-of-concept for GSM use in prosthetic socket development began with CT scan data from (Faustini et al., 2006). The original data is no longer available, but the dimensions and subsequent socket design were used to create a surrogate patient with similar metrics to the subject in this research. Images of the original meshed CT scan elements and their surrogate counterparts are shown in Figure 1.The original patient from (Faustini et al., 2006), was 70 years old, male, 1.82 m tall, and ~90 kg. These same metrics are used in the portion of this research that analyzes socket performance under “patient metrics” with an 800 N force.

These same metrics are used in the portion of this research that analyzes socket performance under “patient metrics” with an 800 N force. A noteworthy difference between the CT scan and this surrogate model is the simplification of the residual skeletal structure as one solid piece. This results in direct translation of forces through the skeletal structure rather than through a layer of soft tissue. While not as accurate to the residual limb anatomy, this decision reduces the problem size, as is required by software used in this research. The lack of soft tissue may decrease the accuracy of the stress locations for some stances; however, it appears to increase the overall magnitude and make the stress in stances with large angles more extreme. With this in mind, the decision is made that this is an acceptable model for this proof-of-concept.A surrogate TSB socket based on the original geometry found in (Faustini et al.

With this in mind, the decision is made that this is an acceptable model for this proof-of-concept.A surrogate TSB socket based on the original geometry found in (Faustini et al. 2006) is developed for the surrogate limb model and serves as a method for validating that a simple socket model based on prior SLS sockets is sufficient for both the patient and a chosen load from the International Organization forFigure 1: comparison of meshed CT scan data (Faustini et al., 2006) and surrogate model developed for this research. (1.1) Original residual skeletal structure, (1.2) simplified solid residual skeletal structure, (2.1) original residual soft tissue meshed from CT scan, (2.2) soft tissue surrogate model, (3.1) original Pe-LiteTM liner meshed from CT scan, (3.2) surrogate liner model.2066Standardization Standard 10328: Prosthetics – Structural testing of lower limb Prostheses (ISO 10328) static proof load 2240 N for case P5 condition 1.

Case P5 refers to a heavier patient of up to 125 kg, and condition 1 refers to loading in the portion of the gait where the reaction force is through the heel rather than the toe. In the simulations performed on the socket in this research, the socket is fixed at the distal end and moments from the ground force on the foot of the prosthesis are not considered. Therefore, the required load for a reaction force more aligned with the distal end of the socket better suits this model. It should be noted that in ISO 10328 an ultimate static test load of 4480 N is included for case P5 condition 1("International Standard (ISO 10328) ", 2016). Due to the nature of this design as a proof-of-concept for GSM in prosthetic socket applications, only the static proof load was targeted for this iteration. The domain for the socket optimization comes from an earlier unsmoothed version of the surrogate socket.

The domain for the socket optimization comes from an earlier unsmoothed version of the surrogate socket. Images of both the smoothed socket, used for validation of the model, and the domain, used for truss development, are shown in Figure 2. The domain is larger than the surrogate model and lacks curvature. This reduces complexity to meet the available problem size for the initial analysis, while increasing the area for the truss model to create novel support geometry.Simulations run on both sockets use the following procedure and boundary conditions. All materials are treated as linear. The distal face of the socket is considered fixed, and all forces are applied to the proximal face of the residual skeletal structure, with the force angle determining the stance. Materials are treated as isotropic; in the case of SLS nylon-12, this means using the material properties in the weakest direction for the entirety of the model.

Materials are treated as isotropic; in the case of SLS nylon-12, this means using the material properties in the weakest direction for the entirety of the model. Mesh resolution is maximized for problem size Figure 2: Left: Smoothed surrogate socket geometry used in FEA, right: constraints in the available software. The surrogate socket is larger unsmoothed bounding box used statically analyzed with both linear and quadratic order tetrahedral as the domain for optimization. meshing for comparison. Linear tetrahedral elements are necessary for the optimization process. However, due to unaccounted for bending of elements, linear order elements tend to raise the safety factor in finite element analysis (FEA). Therefore, a comparison between the two was considered to be a necessary study. The domain is analyzed with linear order tetrahedral elements.

Therefore, a comparison between the two was considered to be a necessary study. The domain is analyzed with linear order tetrahedral elements. Mechanical properties for each material in the model are found in Table 1.Table 1: Material properties for analysisMaterialYoung’s Modulus (Pa) 1.48 e+10Trabecular Bone Limb (soft tissue) Pe-LiteTM Liner SLS Nylon-12 1.9 e+092.5 e+053.8 e+05Poisson’s Ratio 0.300.490.390.39Density (kg/m3) Source(s) and Relevant17641096464.541020Literature (Morgan et al., 2018; Rho et al., 1993) (Faustini et al., 2006; Heymsfield et al., 1989) (Coleman et al., 2004; Rothman, 1962) (ProtoLabs, 2019)The compressive and tensile yield strengths are assigned as 5 e+07 Pa and 7.41 e+08 Pa, respectively, and are found in the datasheet for sintered nylon-12 PA650 supplied by ProtoLabs® (Maple Plain, MN, USA). These yield strengths are typical of the weakest print direction, perpendicular to the bed of the machine (ProtoLabs, 2019).

These yield strengths are typical of the weakest print direction, perpendicular to the bed of the machine (ProtoLabs, 2019). Data from (Hooreweder & Kruth, 2014)2067are used to obtain an approximate S-N curve for sintered nylon-12 in this same direction. Analysis finds both sockets sufficient for supporting residual limb model for the patient load and the ISO 10328 static proof load. The results of the analysis conducted in ANSYS® Mechanical, Release 21.2 (ANSYS, Inc., Canonsburg, PA, USA) are discussed in the findings.In the development of this prosthetic socket, unnecessary material is removed using the ground structure method of topology optimization (GSM). This method is a numerical approximation of an optimized Michell Structure, or truss, where an optimal truss is one with both minimized volume and compliance. While the truss originally described by Michell exists on a continuum, this approximation results in a finite number of trusses.

While the truss originally described by Michell exists on a continuum, this approximation results in a finite number of trusses. Topology optimization using GSM treats an existing geometry, referred to as the domain, or ground structure, as an overconnected truss with members connecting the nodes of the meshed domain space. Members deemed unnecessary are removed to create the least volume, stiffest structure for the given load case (Dorn et al., 1964; Hemp, 1973; Michell, 1904; Prager, 1970).This method is chosen for two primary reasons. First, due to the lack of topological changes to the socket structure over time, there is interest in approaching the loading of a transtibial prosthesis as a problem with structural change inherent to the output. By considering the socket geometry as a redundant truss, the output has already taken a step to further remove itself from increases and decreases in shell wall thickness.

By considering the socket geometry as a redundant truss, the output has already taken a step to further remove itself from increases and decreases in shell wall thickness. Furthermore, it is theorized that if this method is successful, domain spaces further removed from the conventional socket space can be examined to build trusses around the residual limb that more optimally support ambulation or maximize other socket attributes. The second reason for using GSM in this application can be attributed to the extensive body of research reformulating GSM as a linear programming problem, developing closed form solutions, and applying these discoveries to complex domains (Hagishita & Ohsaki, 2009; Hemp, 1973; Rozvany, 1996; Rozvany et al., 1995; Sokół & Rozvany, 2013; Zegard & Paulino, 2014, 2015).

These developments have made GSM an excellent candidate for optimization of a complex loading case on a non-orthogonal structure.Software developed to approach these problems is described in (Zegard & Paulino, 2015). Ground Structure Based Topology Optimization for Arbitrary 3D Domains, or GRAND-3 is a MATLAB® implementation of GSM for domains that have holes and concave regions. This implementation expands on previous work for 2D domains that used a similar linear algebraic approach and video-game development techniques to avoid collision with designated holes that engineers often require for hardware and other components (Zegard & Paulino, 2014). GRAND-3 is based on the formulation of the optimal truss problem as a linear programming problem in (Hemp, 1973). The stiffest truss formulation from (Bendsøe et al., 1994), followed by Hemp’s formulation, is summarized below.

The stiffest truss formulation from (Bendsøe et al., 1994), followed by Hemp’s formulation, is summarized below. Note that in Hemp’s formulation, (Bendsøe et al., 1994)’s notation is still used for continuity.First, examine the loading. Let q and p be vectors of all the nodal forces and the member and is built(axial) forces of a truss, respectively. The nodal equilibrium matrix is defined as BT from the directional cosines of all the members so that:𝑩!𝒒 = 𝒑2068If the volume of each individual bar is represented by ti, the stiffness matrix K of the truss is given by the equation:# 𝑲 = & 𝑡"𝐾" "$%where Ki is the specific stiffness matrix for the ith element. Using K and the displacement vector 𝒖, the optimal truss can be found through:min &,(𝑝!𝒖			 #𝑠. 𝑡.				 & 𝑡"𝐾"𝒖 = 𝑝,"$% # 													& 𝑡" = 𝑉 "$%for	𝑡"	 ≥ 0	,						𝑖 = 	1, . ..

Using K and the displacement vector 𝒖, the optimal truss can be found through:min &,(𝑝!𝒖			 #𝑠. 𝑡.				 & 𝑡"𝐾"𝒖 = 𝑝,"$% # 													& 𝑡" = 𝑉 "$%for	𝑡"	 ≥ 0	,						𝑖 = 	1, . . . , 𝑚If the engineer assigns allowable stresses for the given material, 𝜎! being the allowable tensile stress and 𝜎" being the allowable compressive stress, then 𝑝# 	 ≤ 𝜎!𝑎# and −𝑝# 	 ≤ 𝜎"𝑎# for all values of i.At this point, optimization function is subject to inequalities. The major innovation comes from (Hemp, 1973) adding slack variables developed from p to eliminate these inequalities, creating a linear programming problem. These are defined as:𝑎" =" *𝒊 +#+$ *𝒊 +%& 𝑞" = 𝑠𝒊. - − 𝑠𝒊where 𝑠𝒊% and 𝑠𝒊&are slack variables. This changes the optimization problem to:min /",	/$𝑉 = 𝒍! @𝒔- 𝜎!+𝒔. 𝜎0C𝑠. 𝑡. 				𝑠𝒊≥ 0	,.

- − 𝑠𝒊where 𝑠𝒊% and 𝑠𝒊&are slack variables. This changes the optimization problem to:min /",	/$𝑉 = 𝒍! @𝒔- 𝜎!+𝒔. 𝜎0C𝑠. 𝑡. 				𝑠𝒊≥ 0	,. ≥ 0 𝑠𝒊𝑩!(𝑠𝒊− 𝑠𝒊.) = 𝒒This is the basic formulation for the optimization found in GRAND-3; further informationis available in (Zegard & Paulino, 2015). This formulation and GRAND-3 require certain problem parameters: the deformations must be small, the fixed nodes of the structure must be able to create reaction forces to statically hold the load, and the analysis must be performed for a single load case with linear stress constraints and only elastic deformations.Immediately an issue arises: the analysis of a prosthetic socket is a dynamic problem, typically analyzing the forces as they vary in the walking gait. To address this issue, multiple static analyses are used in socket development.

To address this issue, multiple static analyses are used in socket development. This approach, used by (Faustini et al., 2006), produced a socket that was successfully clinically tested with the patient who is the basis for the surrogate model used in this research. The methods used in this research are similar to quasi- static modeling but have no implied time step, nor do they examine many stances in ambulation. Instead, three stances are selected. These stances represent the peak loads in the sagittal plane:2069heel strike when the foot first descends, standing approximated as vertical, and toe-off when the foot is about to lift into the swing phase of the stance. Freebody diagrams of a full prosthesis in these stances are shown in Figure 3. These diagrams also depict the force Fa applied at the top of the residual skeletal structure in each stance.Figure 3: Freebody Diagrams for a full transtibial prosthesis in three critical positions.

Left: heel strike, center: standing (vertical), right: toe-off.Optimal trusses are calculated for each extreme stance using GRAND-3 in MATLAB® R2022a (Python Software Foundation, https://www.python.org/) implementation is written to create a non-uniform rational b-spline (NURBS) representation of the truss using the same primitives as (Zegard & Paulino, 2015) (cylinders representing the beams and spheres representing the nodes) in Rhinoceros® Version 6 and Version 7 (Robert McNeel & Associates, Seattle, WA, USA).a Python®(MathWorks, Natick, MA USA)andThe trusses generated for the three stances are combined in an attempt to create a low volume stiff truss that tolerates all positions of ambulation. The original output from the GRAND- 3 optimization creates spherical nodes that have the same diameter as the diameter of the member with the largest cross-sectional area.

The original output from the GRAND- 3 optimization creates spherical nodes that have the same diameter as the diameter of the member with the largest cross-sectional area. In the NURBS model generated by the Python® implementation, nodes are generated as spheres with 10% larger radii than the largest member they attach to in order to avoid numerical errors inherent in Boolean operation implementations. Members with a diameter less than 2 mm are increased in size to the 2 mm minimum feature size recommended by ProtoLabs®. Note that SLS with nylon-12 has a minimum feature size of 0.5 mm and can reliably print struts with a 1 mm diameter; however, due to this recommendation and the relatively low mesh density used in this optimization (763 nodes and 2253 elements for the domain), the safety factor provided by the 2 mm minimum diameter members is worth a potential increase beyond the necessary volume for manufacturing.

This method requires the researcher to perform Boolean operations on many components of the truss by hand, as the vast majority of components encounter errors when Boolean operations are attempted in the Python® script. This is due to seam intersection. While NURBS geometry was selected in order to remain true to the geometry, the Boolean process involved slight variations in size and position, deviating from the original truss structure. This Boolean process and the alteration to the geometry is the largest2070disadvantage of using existing NURBS software to create a solid truss. Alternative methods for exporting and attaching GSM geometries more efficiently are discussed in future work.Once connected, the truss system is combined with the critical components at the proximal and distal end of the socket (the brim and the socket base), and post-processing is executed to create a viable socket model.

This process begins with addressing uneven loading of nodes in contact with the residual limb. Loads on the socket interior vary greatly, especially in areas close to the residual skeletal structure. This means that nodes on the socket interior have a range of cross- sectional areas and are not necessarily always in contact with the residual limb as defined in the model. For a successful design, the truss loading needs to be aligned with the static analysis. Furthermore, the primary role of the prosthetic socket is to keep the residual limb stationary relative to the reference frame of the socket. This is a challenge due to the varying node sizes and limited regions of contact. This is addressed by cutting into the nodes with the limb geometry, so the interior evenly contacts the limb. However, this removes volume not only from the nodes but from beams that run along the interior of the socket. Additional volume loss occurs when the file is converted to STL format for printing.

However, this removes volume not only from the nodes but from beams that run along the interior of the socket. Additional volume loss occurs when the file is converted to STL format for printing. The maximum diameter removed from a member due to these changes was determined to be ~2.5 mm. This volume loss is recorded and compensated for in further steps.Due to aforementioned stress variation, members connected to the same node have large deviations in cross sectional area. Since node size is based on the largest beam and is further enlarged for Boolean operations, the smallest beam attached to a node approximates a right angle with the node, creating potential stress concentrations. This issue and volume removal problems are resolved by converting the completed truss structure to a subdivision surface model (SubD). This eases the process of smoothing greatly when compared with other mesh methods.

This eases the process of smoothing greatly when compared with other mesh methods. Several iterations of quadrangulation and smoothing the truss in Rhinoceros®, followed by mesh repair in Materialise Magics® (Materialise NV, Leuvan, Begium), creates a smooth truss model with minimized stress concentrations. The volume loss is compensated for by offsetting the SubD model 1 mm in all directions. Finally, a thin wall is added to the interior creating more points of contact between the socket and the limb and supporting the cut interior nodes and members.

Finally, a thin wall is added to the interior creating more points of contact between the socket and the limb and supporting the cut interior nodes and members. A wall that fully compensates for the maximum volume removal is tested, but it is found that the 1 mm offset combined with the 1 mm wall is sufficient for the loading despite the overlap of truss and wall structures.Once the model is complete, the design is analyzed for both patient and ISO 10328 static proof loads at the three critical stances and at 10% increments between heel strike (0%) and toe- off (100%) to validate that the combined optimization is sound throughout ambulation.

While this does not account for dynamic effects in the ambulation process, there has been success in socket design using this methodology and it is considered sufficient for this proof-of-concept (Faustini et al., 2006; Sotola et al., 2020).FindingsAnalyses of the surrogate limb used with the surrogate socket and domain indicate that the socket is capable of holding the load when fixed at the socket base for angles of heel strike, standing (vertical), and toe-off. This is the case for both the surrogate patient load (800 N) and2071the ISO 10328 static proof load for case P5 condition 1 (2240 N). As a generalization, the domain socket is larger than the surrogate socket model. However, the base of the socket is an exception. It is similar in size and the unsmoothed angles create greater stress concentrations in the area directly proximal to the base. In the final version of this design the brim and base are taken from the surrogate model instead of the domain.

In the final version of this design the brim and base are taken from the surrogate model instead of the domain. While this may protect from stresses in the base, the brim differences have the potential to create inaccuracies within the model. Comparisons of the von-Mises stress of the surrogate socket with quadratic elements, the surrogate socket with linear elements, and the domain are shown in Figure 4 for the ISO 10328 static proof load in the standing (vertical) position.Figure 4: Comparison of von-Mises stress on the surrogate socket with quadratic order tetrahedral elements, the surrogate socket with linear order tetrahedral elements, and the domain socket with linear order tetrahedral elements in the vertical (standing) position at the ISO 10328 static proof load. Shown at 1x deformation.The color scale for each model represents different magnitudes, but stress values in most locations are similar.

Shown at 1x deformation.The color scale for each model represents different magnitudes, but stress values in most locations are similar. The primary difference is that the peak stress on the surrogate model is an area of below average stress on the domain. This is a potential issue in the design process and should be examined in the final model analysis to determine if the socket needs further modification.2072The domain is sufficient for optimization with the surrogate limb and socket model. Due to large forces at the bottom and a lack of deformation in the initial analysis the bottom, 25% of the socket is considered fixed; no truss is created for this area that is replaced by the base from the surrogate model in the final design.

Figure 5 shows the optimal trusses for heel strike, standing, and toe-off, as generated by GRAND-3 and rendered in MATLAB®.Figure 5: GRAND-3 outputs for heel strike, standing, and toe-off with corresponding truss volumes (Zegard & Paulino, 2015).These outputs are imported into Rhinoceros® via Python® with nodes increased by 10% and a minimum cross-sectional area of 2 mm applied to all features. The NURBS truss components modified for assembly and additive manufacturing are shown in Figure 6.Figure 6: Truss models imported into Rhinoceros® adapted for manufacturing and assembly.2073The assembly and post-processing techniques described in the methods section are applied until an acceptable final socket design is created. Images of the truss with the distal and proximal attachments and the final design are shown in Figure 7. The final design is a high-resolution mesh with 691,972 vertices and 1,238,032 faces.

Images of the truss with the distal and proximal attachments and the final design are shown in Figure 7. The final design is a high-resolution mesh with 691,972 vertices and 1,238,032 faces. The final design has a volume of 6.7681 e-04 m3 compared to the original model volume of 8.2963 e-04 m3, 81.58% of the original volume.Figure 7: Final design in Rhinoceros®. Top: Truss with distal and proximal components from the surrogate socket, bottom: complete final design with thin wall.A quadratic tetrahedral mesh with 682,566 nodes and 378,048 is used in the simulations. The thin wall and truss are split into separate components and bonded in ANSYS®. This creates meshable geometries and prevents sliding between the two socket elements in the simulation. The same material properties and surrogate limb model are used. The model is tested for the three critical stances and at 10% increments between heel strike (0%) and toe-off (100%).

The same material properties and surrogate limb model are used. The model is tested for the three critical stances and at 10% increments between heel strike (0%) and toe-off (100%). The safety factors from static analysis and from fatigue analysis (1 e+09 loading cycles) for patient metrics (800 N) and the ISO 10328 static proof load (2240 N) are shown in the charts in Figure 8.2074Figure 8: Safety factors for a single and cyclical load (1 e+09 cycles) of the final design in 10% increments from heel strike to toe-off.The simulation shows that the socket design is acceptable for all stances in static analysis. This is promising but does not necessarily mean the socket will survive dynamic effects. Once fabricated, the socket will be mechanically tested to failure, weak points in the design are noted as they indicate where and how the socket may fail in future testing.

Once fabricated, the socket will be mechanically tested to failure, weak points in the design are noted as they indicate where and how the socket may fail in future testing. The von-Mises stress for heel strike, standing, and toe-off for the ISO 10328 static proof load are shown in Figure 9. Due to the density of meshing on the bars in the truss portion of the design, the simulation results are shown without the elements depicted on the model.2075Figure 9: Simulation results showing von-Mises stress on the final design in heel strike, standing (vertical), and toe-off using the ISO 10328 static proof load (2240 N). Peak and notable stress areas are indicated for each stance. Shown at 1x deformation.In this analysis there are some commonalities in the locations of high von-Mises stress for all three stances, specifically in the anterior lateral corner where the tibia is close to the socket wall and in the posterior of the socket near the proximal end of the solid base.

Both regions should be observed in further testing and may indicate areas for design improvement. In the vertical stance, high von-Mises stress is also noted in the posterior of the socket distal to the brim.DiscussionIn this research a transtibial prosthetic socket is designed using the ground structure method of topology optimization to replace a portion of the socket geometry. This is achieved through the analysis of a surrogate limb and socket model, leading to design decisions to keep the distal and proximal ends of the socket solid. Optimization is run using GRAND-3 (Zegard & Paulino, 2015) on three stances with the surrogate model, each representing peak forces along one of the axes of the sagittal plane. A Python® implementation is written to import the three truss geometries into Rhinoceros® as NURBS primitives where they are assembled into one truss and combined with the distal and proximal components from the surrogate socket.

The model is then post-processed, and the interior is modified to evenly contact the limb geometry. Finally, a 1 mm wall is added to2076increase the contact area between the limb and the socket interior. Static simulations at 10% increments between heel strike and toe-off are run and the design is found to exceed all required metrics at all stances. The thin wall and truss design may reduce the impact of temperature, which is correlated to wall thickness; decrease shear stress through reduction in socket weight; and provide socket wall compliance, mitigating discomfort due to diurnal volumetric fluctuations (Ghoseiri & Safari, 2014; Pirouzi et al., 2014).While this design is successful, there is future work that needs to be done to create a prosthetic socket fit for patient use with this method. Perhaps most critical to the viability of the process is to employ a different approach to the assembly of the truss.

Perhaps most critical to the viability of the process is to employ a different approach to the assembly of the truss. For printing, the truss does not necessarily need to be one solid model. However, in order to run simulations, this is a necessary step. While NURBS models accurately represent the geometry, the process of applying Boolean operations by hand is too time consuming for multiple socket iterations. A mesh output could be used with more advanced software such as Materialise Magics® to correct intersecting facets in an STL representation of truss primitives. A voxel-based model could also be used as Boolean operations are more successful with voxels. While both these outputs change the geometry, similar tactics employed in post-processing could be applied to replace lost volume.

While both these outputs change the geometry, similar tactics employed in post-processing could be applied to replace lost volume. Zegard, the developer of GRAND-3, recommends either using primitives from in X3D format for its simplicity or outputting STL approximations of the tessellated surfaces for GSM export procedure (Zegard & Paulino, 2016).The method for post-processing using a SubD model is successful and is recommended for both its ease and the resulting quality of the final model. Additions to the primitive outputs, for example cone primitives, could be added to the model to decrease the stress concentrations where smaller bars are attached to large nodes. This could further simplify smoothing portion of the post processing procedure.Future work should focus on exploring dynamics.

This could further simplify smoothing portion of the post processing procedure.Future work should focus on exploring dynamics. The results suggest that the socket is adequate for all stances of the walking gait, but without dynamic testing it is unclear if this method produces a socket that is truly optimal for ambulation. The results indicate that applying GSM for static load cases at loading extremes may be a good starting point for the truss of a complex dynamic design problem. Further meshing studies should be conducted to determine the optimal mesh size for analysis and optimization. Furthermore, applying denser meshing in areas of stress concentrations on the surrogate model could produce better results without creating a prohibitively dense mesh for the entirety of the domain.

Furthermore, applying denser meshing in areas of stress concentrations on the surrogate model could produce better results without creating a prohibitively dense mesh for the entirety of the domain. Different domain spaces could also be used to further remove material from the original geometry and determine if a different over all structure could be used, e.g., a truss that supports a thin wall entirely from the lateral side of the limb to avoid interference with the user’s other leg. Nevertheless, this first proof-of-concept illustrates that socket geometry can be adapted and optimized. Further exploration into optimal structures for stiffness, patient comfort, and other target metrics can and should be conducted to determine how AM and topology optimization can impact the stagnant design of transtibial prosthetic sockets.2077ReferencesBallit, A., Mougharbel, I., Ghaziri, H., & Dao, T.-T. (2022).

SIGN AND ANALYSIS OF ORTHOGONALLY COMPLIANT FEATURES FOR DURAFORM/SLS MANUFACTURED PLATESMario Faustini*, Richard Crawford*, Richard R. Neptune*, William Rogers†, Andrew Gitter† and Gordon Bosker†The University of Texas at Austin †The University of Texas Health Science Center in San AntonioAbstractIn many applications of parts manufactured by Solid Freeform Fabrication (SFF), compliance is an important factor. In order to achieve given deformation goals with optimal shape, the design of compliant mechanisms and elements fabricated with SFF techniques must take into account the particular constraints and boundary conditions of the target application as well as the specific material properties of the part. The present work focuses on the design and evaluation of compliant features for a geometrically constrained thin-wall part subject to loads normal to its tangent plane. Such features would need to be embedded in the object.

Such features would need to be embedded in the object. The manufacture of prosthetic sockets for lower-limb amputees is the specific application presented, where greater compliance is needed at sites in contact with pressure sensitive tissues. Sample parts were fabricated by selective laser sintering, and the material used was Duraform.1IntroductionCompliance is in many cases a fundamental and desired property of a mechanical structure. In many applications, compliance is very important for proper function of a part. In others it is a much more elegant and functional design solution that substitutes joints and links of mechanisms by incorporating controlled flexibility within the structure itself. In order to proceed with the design of compliant features, very strict constraints must be considered and evaluated [Howell, 2001].

In order to proceed with the design of compliant features, very strict constraints must be considered and evaluated [Howell, 2001]. The part material and its mechanical properties must be studied, boundary conditions for the structure must be well defined, the target deformation under specific loads must be set, and a thorough failure prevention study, using both static and fatigue theory, should be performed.This article focuses on the primary design stages of orthogonally compliant features of plates and thin walls (where the load/displacement direction of the feature is normal to the plate or wall), and evaluating the design options that are more promising for a given set of design goals for an specific application.

The design example to be considered is the incorporation of compliance features in sockets for below-the-knee prostheses, produced with selective laser sintering (SLS) using Duraform.2Fabrication of Sockets for Below-the-Knee Prostheses using SLSThe production of sockets for prostheses using SFF (Figure 1) has several potential advantages compared to common, more labor intensive and time-consuming fabrication methods [Walsh et al., 1989; Freeman et al., 1998]. One of the main advantages is that SFF directly creates sockets400from digital shape information, eliminating the need for molds, hand lamination and finishing procedures.

One of the main advantages is that SFF directly creates sockets400from digital shape information, eliminating the need for molds, hand lamination and finishing procedures. Another advantage of SLS is the ability to create complex geometries with minimal cost penalty in manufacturing, which significantly expands the options for developing and exploring alternate socket designs.Figure 1: On the left and center, a patient testing an SLS fabricated socket; on the right, the description of the parts of a prosthesis for below-the-knee amputees.Our method, developed through collaboration of the Laboratory for Freeform Fabrication at The University of Texas at Austin and the Rehabilitation Engineering Laboratory of The University of Texas Health Science Center at San Antonio [Rogers et al., 2001], creates a computer model of the socket from digital source taken from the patient’s residual limb.

The shape of the socket is designed so that the load is supported by specific regions of the limb that can safely resist contact pressures generated when the patient walks. This design is called the Patellar-Tendon Bearing (PTB) socket, illustrated schematically in Figure 2.Some areas of the residual limb are not suitable for high contact pressures, either because they are too sensitive to pain, are susceptible to developing sores, or other various health issues. Notwithstanding the reason, high pressures at these sites lead to discomfort for the patient during gait (thus affecting rehabilitation) and, sometimes, pose a health risk for the patient. Thus, it is necessary to incorporate means of relieving contact pressure over these areas. One way to achieve this is to make the socket orthogonally compliant at these high pressure contact sites.Currently, local compliance is achieved by reducing the wall thickness of the socket at pressure sensitive areas, as seen in Figure 3.

However, this approach provides limited compliance and pressure relief. A more effective design may be to take advantage of the ability of SLS to easily create geometric variants of traditional socket shapes that allows a controlled expansion of socket volume to adapt to limb change during gait.401Figure 2: Schematic of a socket based on the PTB approach, with main force vectors shown.Figure 3: Example of a current SLS fabricated socket with locally thinned walls in areas where compliance is needed.3Design Goals and ConstraintsIn order to allow the effective and successful design of orthogonally compliant features to be incorporated into the socket structure, it is necessary to identify the design goals and constraints. First, a target deformation for the compliant region when subjected to the normal steady-load pressure must be specified.

First, a target deformation for the compliant region when subjected to the normal steady-load pressure must be specified. The optimum value of the desired displacement at the center of the compliant site is, however, difficult to set precisely, since it depends on several parameters that may vary from patient to patient and can be highly subjective. Our experience has shown that positive results occur when the target displacement should be between 2 and 4 mm.402Second, the incorporated compliant features on the wall of sockets should not compromise structural safety during gait. Moreover, such features should not affect the shape of the inside of the socket and should have minimal impact on its outer appearance. The latter condition, although very subjective, is nevertheless important for the patient to feel positive about using the prosthesis.Third, areas with no material on the socket wall, such as slots and holes, should have minimal dimensions.

Since the limb requires a constant minimal pressure to be applied to all areas, even on more sensitive areas, large holes on the wall should be avoided altogether. Features that lead to high localized contact pressure gradients and concentrations should also be avoided, since they can lead to circulatory problems for the patient.Finally, compliant features should not produce spikes or sharp corners when deformed. Hence, slots on the socket wall should not self-intersect or cross each other. Figure 4 shows an example of a possible socket with compliant features that satisfy these design goals and constraints.Figure 4: Example of orthogonally compliant features incorporated in a socket.4FEM Evaluation of Some Design OptionsFor the evaluation of various design options that satisfy the design constraints described in the previous section, computer modeled test discs were evaluated using finite element analysis (FEA) and prototypes were produced with SLS and Duraform material.

These discs were dimensionally similar to compliant areas on the socket, as shown in Figure 5. Geometric features such as slots and holes will be added to each disc to assess their ability to provide compliance.403Figure 5: Test disc to evaluate compliance options (dimensions in meters)The boundary conditions for the FEA analysis where defined as follows: the borders of the discs were fully restrained and a pressure of 10 kPa is applied at one of the free faces. SDRC I-DEAS 8 was used for the analysis and all elements were defined as 10-node parabolic tetrahedra. The material properties used for Duraform were those published by 3D Systems (Valencia, CA).Figure 6 shows the deformation results for a solid control disc, with no compliant features. The displacement at the center due to the 10 kPa pressure was 5.54 x 10-5 m. Figures 7, 8 and 9 show the results for 4, 8 and 16 straight slots radially distributed on the discs, respectively.

The displacement at the center due to the 10 kPa pressure was 5.54 x 10-5 m. Figures 7, 8 and 9 show the results for 4, 8 and 16 straight slots radially distributed on the discs, respectively. The central displacement achieved varied from 6.27 x 10-5 to 7.20 x 10-5m as the number of straight slots was increased. However, a design with more than 16 slots would start to compromise structural integrity due to higher stress concentrations.404Figure 6: FEA deformation results for control disc with no special compliant features.405Figure 7: FEA deformation results for test disc with 4 radial slots.406Figure 8: FEA deformation results for test disc with 8 radial slots.407Figure 9: FEA deformation results for test disc with 16 radial slots.408Figure 10 presents an option with 4 larger void spaces, which is topologically similar to the solution given by optimization methods.

For this design, the central displacement reaches a value of 8.52 x 10-5m, which is still not as high as desired. In addition, there is a potential problem with localized pressure gradients (due to contact discontinuity) at the borders of the holes that may affect the peripheral circulation in the residual limb.Figure 10: FEA deformation results for test disc with 4 void spaces.409Finally, Figures 11 and 12 display the results for designs with spiral slots (2 and 4 radially distributed spirals, respectively). For the case with 2 spirals, a central displacement of 4.38 x 10-3 m was achieved, which is very satisfactory.

For the case with 2 spirals, a central displacement of 4.38 x 10-3 m was achieved, which is very satisfactory. The case with 4 spirals, on the other hand, results in a central displacement of 8.88 x 10-3 m, which was considered to be too large.Figure 11: FEA deformation results for test disc with 2 spiral slots.410Figure 12: FEA deformation results for test disc with 4 spiral slots.4115Conclusions and Future WorkThis article presents a preliminary evaluation of design options for orthogonally compliant features to be used in sockets for below-knee prostheses using SLS. Of the options investigated, the most effective design incorporated two spiral slots, which balanced providing the needed displacement while avoiding high contact pressures.This work is a preliminary study of specific topologies for compliant features using a trial and error approach.

Future work will be directed at applying a modified version of topology optimization method to systematically optimize designs that satisfy the design constraints. One such technique is the Homogenization Method [Bendoe et al., 1993; Suziki and Kikuchi, 1991; Nishiwaki et al., 1998], which considers a design domain described by a 2D microstructure of unit cells, each one with void space (with sides defined by values a and b, which vary from 0 – no material – to 1) and orientation (defined by angle q) (Figure 13).

In general terms, the method consists of a recursive approach that calculates homogenized elasticity coefficients for the design domain, calculates sensitivities with FEM, optimizes for compliance with respect to a and b, filters results to eliminate checkboard patterns, updates angle q and repeats all steps until convergence is achieved.abqUnit cellDesign domain: microstructureFigure 13: Design domain for the Homogenization Method.However, some modifications must be applied to this method to allow its use with orthogonally compliant plates. First, a microstructure with orthogonal boundary conditions must be developed. Another fundamental modification is to filter solutions with large void areas, which violate design constraints for socket application.

Another fundamental modification is to filter solutions with large void areas, which violate design constraints for socket application. Finally, fatigue analysis should be performed on the compliant features.412AcknowledgementsThe authors gratefully acknowledge financial support for this work from the VA Rehabilitation Research and Development Service.ReferencesBendoe, M. P., Diaz, A., and Kikuchi, N., 1993, “Generating Optimal Topologies in Structural Design Using a Homogenization Method,” Computer Methods in Applied Mechanics and Engineering, 71 (1988), pp. 197-224Freeman, D., Wontorcik, L., 1998, “Stereolithography and Prosthetic Test Socket Manufacture: A Cost/Benefit Analysis”, Journal of Prosthetics and Orthotics, Vol. 10, Num. 1, 1998.Howell, L. L., 2001, Compliant Mechanisms, John Wiley & Sons, New York.Nishiwaki, S., Frecker, M., Min, S.

SIGN AND ASSESSMENT OF AN AM VENDING MACHINE FOR STUDENT USENicholas A. Meisel and Christopher B. Williams Design, Research, and Education for Additive Manufacturing Systems Laboratory Department of Mechanical Engineering & Department of Engineering Education Virginia TechABSTRACTDue to prohibitive costs, access to Additive Manufacturing (AM) technologies at academic institutions tends to be limited to upper-level courses that feature significant project-based coursework, such as capstone design. However, with the decreasing cost of desktop-scale AM technology, there is potential to improve student access to such technologies throughout a student’s undergraduate career, and thus provide more opportunities for AM education. In this poster, the authors present the design and implementation of an AM “vending machine” that is powered by desktop-scale extrusion-based AM systems.

In this poster, the authors present the design and implementation of an AM “vending machine” that is powered by desktop-scale extrusion-based AM systems. The resulting machine allows for unrestricted student use of AM equipment, and thus provides ample opportunity for informal learning regarding AM. The results of a formal assessment of student use of the machine are presented.Keywords: 3D Printing, Material Extrusion, Vending Machine, Informal Learning1. STUDENT EXPOSURE TO MANUFACTURINGIn the “2009 Roadmap for Additive Manufacturing,” the workshop participants identified education as a critical area of research and development for the advancement of the Additive Manufacturing (AM) field [1]. Fundamentally, it was thought that current and future engineers’ familiarity with AM technologies was lacking, and as a result, was a barrier to AM adoption.

Fundamentally, it was thought that current and future engineers’ familiarity with AM technologies was lacking, and as a result, was a barrier to AM adoption. As a result, the document urged the development of university “programs for educating the general population to enhance the interest in AM applications and generate some societal ‘pull’ for the technologies” [1]. AM curricula is emerging in U.S academic institutions (e.g., [2, 3]), but is not yet pervasive.learning experiences with AM, and advanced manufacturing technologies in general, has historically been cost prohibitive for most academic institutions. Specifically, costs of purchasing, operating, and maintaining the manufacturing equipment (and costs related to staffing) have often limited the number of institutions with interactive manufacturing curricula. As a result, student access to manufacturing technologies is typically limited to classroom and class laboratory settings.

As a result, student access to manufacturing technologies is typically limited to classroom and class laboratory settings. While these contexts provide valuable experiences, they are inherently constrained by the classroom itself. Students have limited hours to access the equipment, have limited amount of working material, and most importantly, are working within the confines and rule sets of an assignment. Unfortunately, these limitations somewhat constrain student learning. Emerging engineering education research suggests thatProviding hands-ondesign-based, semi-autonomous experiences in informal learning settings provide important opportunities for self-directed learning.1.1. Manufacturing Informal Learning Environments: Opportunities and ChallengesMost of educational pedagogical practice, development, and research is focused in crafting meaningful experiences in formal learning environments; i.e., learning within the confines of a classroom.

However, there is increasing interest in informal learning environments – extracurricular activities such as internships, design competitions, and student engineering professional societies – as they have been shown to positively contribute to students’ engineering education [4]. These informal learning environments are characterized as unstructured, socially- based educational environments wherein students collaborate autonomously on a project. A hallmark of these environments is that they are “person centered,” meaning that the objectives and goals of the activities are selected by the students, and not imposed by an instructor.Research of informal learning environments suggests that they can promote and enhance student skills in leadership, collaboration, communication, innovation, and motivation [5]. In addition, informal learning environments have shown to be effective for promoting learning outcomes related to interdisciplinarity [6] and design education [7].

In addition, informal learning environments have shown to be effective for promoting learning outcomes related to interdisciplinarity [6] and design education [7]. These results are supported by the situated cognition learning theory, which states that knowledge is contextual and is a product of the situations in which it is created, and is thus positively influenced by professionally authentic activities [8].Manufacturing-related informal learning environments would resemble student machine shops (or equivalent spaces) wherein students are provided autonomous control of the facility to explore manufacturing in a project-based context that is relevant to a personal interest. However, due to the aforementioned resource constraints, few institutions offer such opportunities.

However, due to the aforementioned resource constraints, few institutions offer such opportunities. Those that exist (e.g., Virginia Tech’s Ware Lab [9], Penn State’s Learning Factory [10, 11], Georgia Tech’s Invention Studio [12] and emerging Makerspaces and Techshop) are often constrained by factors such as:Scale: The labs are often not large enough to support the entire student body interested in making.Safety and Insurance: Due to safety concerns of the students (and the equipment), institutions are often relunctant to provide true open access to manufacturing spaces. To overcome this requirement, spaces often require significant training (taking the form of formal classes in some cases), staff oversight, and insurance costs.

To overcome this requirement, spaces often require significant training (taking the form of formal classes in some cases), staff oversight, and insurance costs. Institutional barriers to access: In addition to potential pre-requisite safety classes for shop use, most spaces employ institutional barriers to access such as limits to students of certain departments/colleges and/or of specific design projects (e.g., competition teams). Furthermore, it is rare that shop space is available to students who wish to work on projects that are not directly related to a specific class or research project. This is often justified by the desire to spend a department’s resources on only its students.While these constraints provide a structured way of managing institutional resources, they unfortunately impose inexplicit barriers to student access to informal manufacturing learning opportunities.1.2.

ContextRecognizing the importance of exposing future engineers to AM technologies, the authors seek to provide an open-access environment in which students can experience AM (and manufacturing in general). This vision is motivated by the constraints of the existing manufacturing-related learning environments at Virginia Tech and similar large-scale academic institutions. The Virginia Tech College of Engineering offers several AM-related student experiences: the Frith Lab features a Stratasys uPrint system for a few selected design teams; the staff-run machine shops within the departments of Mechanical Engineering and Aerospace Engineering offer AM as an option for realizing students’ classroom design projects; the DREAMS Lab, a AM research lab, offers limited support for using AM for specific design and research projects; and there exists a dedicated AM course for senior undergraduate students and graduate students in the College of Engineering.

Despite these opportunities, student-access to AM is limited, especially given the scale of the Engineering student body (8,642 students).In the face of these limitations, the authors desired to develop a mechanism for increasing student access to manufacturing tools and to increase student awareness of modern manufacturing challenges and considerations. Towards this goal, the authors have created an AM “vending machine” for use by all faculty, staff, and students. The purpose of this paper is to document and share the design rationale, the implementation details, and assessment of the system in hopes that these details might aid others’ future embodiments. The authors present an overview of the overall system design goals and rationale in Section 2. A discussion of the system embodiment and implementation is offered in Section 3.

The authors present an overview of the overall system design goals and rationale in Section 2. A discussion of the system embodiment and implementation is offered in Section 3. Preliminary assessmet of the impacts of the system on student learning is provided in Section 4, including both qualitative (Section 4.1) and quantitative (Section 4.2) results. Closure and thoughts on future work are offered in Section 5.2. OVERVIEW OF THE AM VENDING MACHINE SYSTEMRecognizing the challenges in offering hands-on manufacturing educational experiences at a large-scale institution such as Virginia Tech (Section 1.1), the authors were driven to create a means for students to experience AM technologies with almost no barrier to entry. From this overall goal, the team conceived of an AM “vending machine” system concept. The overall concept of an AM “vending machine” is a system that automatically fabricates and dispenses parts on-demand at the users’ request.

The overall concept of an AM “vending machine” is a system that automatically fabricates and dispenses parts on-demand at the users’ request. Instead of providing money in exchange for a fixed piece of inventory, a user provides an AM vending machine a digital part file of his/her own design.The vending machine interface was identified as a novel means of providing students access to AM, that also circumvented inherent limitations of formal experiences, for the following reasons:The system provides means of educating students about AM via hand-on interaction with the technology and via a display of parts fabricated via a wide variety of AM systems. Furthermore, a vending machine interface provides a means of educating students aboutAM at a large scale: no registration or formal training is needed to use the system, and it is available at all times.The system provides a user interface that is neither intimidating nor imposes perceived barriers to entry.

Unlike a dedicated room or laboratory, which might feature locked doors or require user accounts, a vending machine has an interface that is immediately comfortable to the user and imposes no preconceived notion of ownership. Traditional vending machines are operated by everyone and do not require “permission” to use.  The system can be placed in a public space, such as a hallway, and thus have high visibility and high student interaction.The system can be designed and operated economically, and thus can be used by students to create parts for any purpose: design project, classroom assignment, personal projects, etc.A vending machine system illustrates the ability of AM to provide just-in-time manufacturing solutions. In effect, interface provides a juxtaposition: while traditional vending machines feature a fixed inventory, a vending machine powered by AM systems provides an infinite inventory by delivering user- defined products on demand. the vending machine3.

the vending machine3. AM VENDING MACHINE SYSTEM EMBODIMENTFrom the overall vision and design goals for an AM vending machine (Section 2), the team created “The DreamVendor,” a vending machine system powered by four MakerBot Thing-O- Matic (TOM) desktop-scale extrusion systems [15], as shown in Figure 1. The system case is composed of an extruded aluminum frame, laser-cut acrylic, and custom-made components made via laser sintering (e.g., bin handles, bin hinges, Interface Board mounts).(a)(b)Figure 1. Images of (a) the DreamVendor system and (b) the Thing-o-Matic extrusion machines in operation.A student begins by loading his/her design (in the form of the MakerBot .s3g toolpath file) onto a SD card. The student loads the SD card into the machine's control panel, and selects the file to print. Once the print is complete, the part is ejected into a bin to be retrieved.3.1.

The student loads the SD card into the machine's control panel, and selects the file to print. Once the print is complete, the part is ejected into a bin to be retrieved.3.1. System Implementation DetailsAt an abstract level, an AM vending machine system must perform four primary functions: (i) receive user input, (ii) fabricate part, (iii) dispense part, and (iv) assess the user experience. In this sub-section, the authors detail the manner in which the DreamVendor accomplishes each of these sub-functions.Receive User InputThe primary “vending machine” user interface is driven by MakerBot Interface Boards [16] that are connected to each of the system’s TOMs (Figure 2a). The interface boards are installed onto the face of the display cabinet using custom designed holders that were manufactured via laser sintering. Using their own personal computers, students generate G-code from their designed .STL file using ReplicatorG software [17].

Using their own personal computers, students generate G-code from their designed .STL file using ReplicatorG software [17]. Students’ file preparation is guided by thorough instructions provided at the DreamVendor website [18]. The resulting G-code is then transferred to a SD card before coming to use the DreamVendor. Once they arrive, the users input their SD card into the appropriate slot on the Interface Board. Using the buttons on the board, they can then view a list of files and select the appropriate file for fabrication (Figure 2b).(a)(b)Figure 2. (a) The DreamVendor cabinet and (b) user interface board and part ejection bin.The team considered many other potential solutions for the “receive user input” function. Initial concepts featured users bringing STL files to the vending machine via memory stick; however, it was determined that it was not ideal to have users waiting to slice and generate toolpath while standing at the machine.

“Cloud-based” submission solutions (i.e., a cyber- enabled system that allows users to submit files over the internet) were also considered; however, they were not selected for their implementation difficulty.Fabricate PartOnce a part’s G-code is selected on the user interface board, the TOM first preheats the build platform and nozzle, and then begins fabricating the part in a layerwise fashion. The modified Interface Board firmware displays the print time and percent completion during a build on the LCD screen.To ensure that the provided files are compatible with the DreamVendor TOM hardware, the team created a unique “DreamVendor” ReplicatorG profile that students are instructed to use to prepare .STL files. In addition, the Interface Boards’ firmware was modified to include a check of the user-provided G-code; if it is detected that the user failed to use the specified profile, the machine will not print the selected part.

These precautions are necessary to ensure that users do not accidentally provide G-code that could harm the machines. Finally, the modified firmware provides a secret maintenance menu (accessed by a sequence of buttons on the Interface Boards) that allows users to input basic maintenance commands (e.g., jogging the motors, heating or cooling the printheads, etc.).Dispense PartOnce a user’s part has finished printing, a system is necessary to remove the part from the build platform so that other users may use the machine. To this end, the Thing-O-Matic’s Automated Build Platform (ABP) was used for ejection. The ABP is composed of a conveyor belt that wraps around a TOM’s heated build platform. When driven by its motor at the conclusion of the build, the conveyor belt effectively removes the part from the system as it conveys across the front edge of the build platform. Upon ejection, the part lands in a part collection bin that is mounted into the face of the cabinet.

Upon ejection, the part lands in a part collection bin that is mounted into the face of the cabinet. The bin and the interface board are shown in Figure 2b. To ensure that all new builds begin on an empty ABP, the DreamVendor ReplicatorG profile has been modified to include a command in the G-code to turn the conveyor belt before each print.Assess the User ExperienceOne drawback of using SD cards containing G-code as the primary source of user input is that there does not exist a manner in which to capture data on number and content of printed parts, which would be a valuable resource for evaluating the efficacy of the DreamVendor. Due to this limitation, a secondary assessment instrument was devised to collect information from users.Specifically, an electronic survey kiosk has been installed onto the side of the vending machine. This system, powered by an Android tablet, allows for real-time assessment and feedback of student use of the DreamVendor.

This system, powered by an Android tablet, allows for real-time assessment and feedback of student use of the DreamVendor. The assessment instrument, an eight-question survey, is driven by a custom-made DreamVendor application. The results from this survey, exported as a .TXT file, are automatically uploaded to a DropBox folder each evening via the tablet’s embedded wifi. Results from this assessment are discussed in Section 4.3.2. System Operation DetailsThe DreamVendor was designed, manufactured and installed by a volunteer team of four graduate students over the course of one year. Initial funding for the system was provided by a student organization; continued financial support, which is used to purchase raw material for the TOMs, is provided by student enrollment fees.

Initial funding for the system was provided by a student organization; continued financial support, which is used to purchase raw material for the TOMs, is provided by student enrollment fees. Due to these contributions, the system is completely free to use for all faculty, staff, and students of the university, regardless of their year or departmental affiliation.The DreamVendor was installed in the main lobby of the Randolph Hall at Virginia Tech, which is home to the Aerospace and Mechanical Engineering departments, as well as classrooms for first-year engineering curricula. This location ensures tremendous foot-traffic by the DreamVendor, and provides a broad exposure.As the TOMs are of a hobbyist, desktop scale, they are not engineered for around-the-clock fabrication that is required for the “vending machine” context. As such, system maintenance is often a continuous process.

As such, system maintenance is often a continuous process. The DreamVendor’s inclusion of four TOMs provides a modular design; if one is under maintenance, it can be either replaced by another system, or repaired without disrupting the other three functional systems. Day-to-day maintenance, which can include unclogging nozzle heads, reloading material, clearing crashed builds, and lubricating machine axes, is conducted by a team of undergraduate student volunteers.4. PRELIMINARY ASSESSMENT OF STUDENT USEOriginally unveiled in March 2012, the DreamVendor system has become a focal point for the Virginia Tech undergraduate engineering curriculum. It has also become a central point of prospective student tours and outreach activities (Figure 3). This outcome is due to both the DreamVendor’s role as a prototyping resource for all students and its nature as a catalyst in creating an informal learning environment.

This outcome is due to both the DreamVendor’s role as a prototyping resource for all students and its nature as a catalyst in creating an informal learning environment. Preliminary assessment of the system is provided in this section using both qualitative (Section 4.1) and quantitative (Section 4.2) sources.Figure 3. An outreach event targeting under-represented groups visits the DreamVendor4.1. Qualitative AssessmentSince its unveiling, the DreamVendor system has been in almost constant use during its 8hours of open operation each day. Student use can be classified into three categories:Experiential parts: Students initial experience with the DreamVendor tends to be focused in participating in the experience.

Student use can be classified into three categories:Experiential parts: Students initial experience with the DreamVendor tends to be focused in participating in the experience. As such, parts are often tokens, statuettes, and other aesthetic tchotchkes that have been downloaded from the internet.Prototype parts: Providing open access to AM has provided many students the resource needed to create prototypes of their novel designs for personal and classroom projects. One example is a design team, which printed models of fuel-cells for planning the layout of their hybrid-electric competition vehicle. In addition, students conducting analytical research have also used the system to print physical prototypes of the systems they are virtually analyzing.Functional parts: Students have also used the DreamVendor to create parts that they implement and use. For example, students from Design/Build/Fly competition have used the AM vending machine to fabricate bay doors for their aircraft.

For example, students from Design/Build/Fly competition have used the AM vending machine to fabricate bay doors for their aircraft. In addition, students have created parts for their personal projects, such as the (i) cellphone solar charging device, (ii) monoscopic camera/mirror mount, and (iii) custom toothbrush holder shown in Figure 4.(a)(b)(c) Figure 4. Functional objects created on the DreamVendor: (a) solar-powered cellphone charger, (b) monoscopic camera/mirror mount and (c) custom toothbrush holder.Beyond its role as a formal student prototyping resource, the DreamVendor system has also exhibited a significant informal learning component. It is not uncommon during the day to see students crowd around the DreamVendor while it is in operation to discuss concepts related toCAD/CAM, AM, design, and manufacturing. In fact, students have reported learning CAD software just to be able to use the machine.

In fact, students have reported learning CAD software just to be able to use the machine. The DreamVendor’s undergraduate maintenance team members also have the opportunity to gain hands-on experience in a basic research capacity, coming up with ways to further increase the impact of the DreamVendor. In addition, an extracurricular design competition (“CreatiVT”) that awards the “Best Use of the DreamVendor Award” is held each semester to help encourage student design thinking and to help build a community of use.4.2. Quantitative AssessmentAs noted in Section 3.1, the team’s ability to formally assess student use of the DreamVendor is limited by the fact that the system does not retain copies of users’ prints. To provide some means of assessing student impact, the team installed a survey kiosk onto the system. The custom Interface Board firmware asks students to complete the survey before every build.

To provide some means of assessing student impact, the team installed a survey kiosk onto the system. The custom Interface Board firmware asks students to complete the survey before every build. This survey is voluntary and seeks to collect basic information on the demographics of the students using the system as well as on the parts that they are printing. Over the course of 5 months of operation, the survey has had 186 respondents.For demographic information, students are asked which college they belong to within Virginia Tech, which department they belong to, and what year of study they are currently in. Results for each of these questions are presented in Figures 5, 6, and 7, respectively.Figure 5. College affiliation of DreamVendor usersFigure 6. Departmental affiliation of DreamVendor usersFigure 7. Year of study of DreamVendor usersFrom this data, it is clear that the majority of users are associated with the College of Engineerin.

Departmental affiliation of DreamVendor usersFigure 7. Year of study of DreamVendor usersFrom this data, it is clear that the majority of users are associated with the College of Engineerin. While other university colleges are represented, it is in small proportion to theCollege of Engineering. Within the college of Engineering, the majority of users associate with mechanical engineering (60% of sample), with a significant number also associated with Aerospace Engineering (14% of sample). This result aligns well with the field of AM, which is most likely to resonate with Mechanical and Aerospace Engineers at this point in their academic careers. It also aligns well with the DreamVendor’s location in the building which houses both departments.The year of study distribution appears to be more evenly divided among the users, with Freshman students in the majority (26% of sample) and Seniors in second by a small margin (25% of sample).

This data suggests that the goal of offering prototyping capabilities to students involved in hands-on design coursework (cornerstone design for freshmen, capstone design for seniors) has been achieved.Data collected regarding student part creation is similarly illuminating. Students were asked, in general terms, why they were printing the part, whether or not the part was decorative or functional, and how they acquired the part design. Results from the first of these questions (Figure 8) clearly shows that the majority of the students are creating parts for personal use (73% of sample). A sizable portion of students also created parts in fulfillment of their coursework (12% of sample). Student responses to the next question (decorative vs. functional) shows that 50% of users are creating functional parts while 50% are creating decorative parts.

Student responses to the next question (decorative vs. functional) shows that 50% of users are creating functional parts while 50% are creating decorative parts. This distribution suggests that a sizable portion of student users see the potential in AM processes, even simple desktop-scale extrusion, to create functional, end-use parts. The final question on the survey, asking students how the acquired the part, provided incredibly promising results (Figure 9). 72% of parts made on the DreamVendor system were designed and created by the students themselves. Only 25% of parts were downloaded from online websites such as Thingiverse. This result demonstrates that the majority of students are using the system to create something of their own design, rather than simply acting as consumers of someone else’s design. This helps validate the DreamVendor’s position as an effective tool to encourage student learning.Figure 8. The intended use of parts created using the DreamVendorFigure 9.

This helps validate the DreamVendor’s position as an effective tool to encourage student learning.Figure 8. The intended use of parts created using the DreamVendorFigure 9. Acquisition source of parts created on the DreamVendorIn addition to the quantitative data presented above, students were asked to provide anecdotal evidence of their DreamVendor by providing more specific details about their printed part as well as leave any additional comments about the DreamVendor system. Student comments reveal a wide variety of printed parts for a wide range of uses. Printed parts include a model rocket body, an automatic fish feeder for while a student was away on vacation, various phone stands and cases, as well as a horse massager designed and created by a Pre-Veterinary student.

Additional student comments include,“Thank you for having these for free! I love being able to use them and I want to learn CAD more because of this and I have actually started to learn inventor more because of this!”;“This is awesome! Thank you for letting students use this facility.”;  “Great opportunity for all students, I am glad it is free.”While these comments only represent the thoughts of a portion of all users (most users declined to provide additional comments), they clearly demonstrate student appreciation for the manufacturing opportunities afforded to them by the DreamVendor System.5. CLOSURE AND FUTURE WORKIn this paper, the authors have discussed the design and implementation of the DreamVendor public use AM system. The overall design of the system is discussed, as well as specific methods for addressing the user interface, part ejection, assessment, system maintenance.

The overall design of the system is discussed, as well as specific methods for addressing the user interface, part ejection, assessment, system maintenance. Preliminary assessment has shown that the system meets the goal of encouraging students of all grade levels to learn more about AM and CAD and to create their own parts. Of all the educational benefits offered by the DreamVendor system, the informal learning component is the most surprising.Future work will focus on improving both the reliability of the system as well as its portability. In addition, assessment will be expanded to gain more detailed quantitative and anecdotal evidence regarding how students are using the DreamVendor to expand their knowledge of manufacturing.6. ACKNOWLEDGEMENTSThe authors acknowledge the tireless efforts of the many students who volunteered their time to make the DreamVendor a reality.

ACKNOWLEDGEMENTSThe authors acknowledge the tireless efforts of the many students who volunteered their time to make the DreamVendor a reality. The team that led the creation of the DreamVendor was composed of four graduate students from the Department of Mechanical Engineering: Amelia Elliott (lead designer), David McCarthy, Jacob Moore, and Nicholas Meisel. Since its inception, the DreamVendor has been supported by a diligent team of students from across the College of Engineering: Andrew Price, Kevin Kline, Corey Buttel, Hannah Thomas, Todd Spurgeon, Matthew Price, Matt Ocheltree, Callie Zawaski, Brandon Hart, and Josh Eddy. The authors also acknowledge the financial support of the Virginia Tech Student Engineers’ Council, the Department of Mechanical Engineering, and General Motors.7. REFERENCES[1] Bourell, D., M. C. Leu and D. W.

sign and characterization of orthotropic re-entrant auxetic structures made via EBM using Ti6Al4V and pure copper L. Yang*, O. Harrysson*, H. West II* and D. Cormier†Edward P. Fitts Department of Industrial and Systems Engineering, North Carolina State University, Raleigh, NC 27606 †Department of Industrial and Systems Engineering, Rochester Institute of Technology, Rochester, NY 14623AbstractAn orthotropic 3D re-entrant honeycomb structure that exhibits a negative Poisson’s ratiowas designed and fabricated via the electron beam melting (EBM) process. The modeling work established the relationships between various structural parameters and the mechanical properties of the auxetic structures. Compressive tests were performed on the re-entrant honeycomb samples made with Ti6Al4V as well as pure copper.

Compressive tests were performed on the re-entrant honeycomb samples made with Ti6Al4V as well as pure copper. Results of the strength, modulus and energy absorption for the two materials were compared with the theoretical models in order to verify the theoretical predictions.Keywords: Electron Beam Melting, Auxetic Structures, Ti6Al4V, copperIntroductionAuxetic structures have shown promise for various engineering applications such assandwich panel cores, impact absorption, biomedical implants and aerospace structures. These structures exhibit high shear modulus [1-3], high indentation resistance, high fracture toughness [4-6], high energy absorption ability [7-10], and novel acoustic and electromagnetic absorption abilities [12-14]. The prevailing manufacturing process for synthesis of polymer auxetic structures has somewhat limited control over the mechanical properties [1, 15].

The prevailing manufacturing process for synthesis of polymer auxetic structures has somewhat limited control over the mechanical properties [1, 15]. As a result, much of the modeling work involving auxetic structures has lacked experimental validation, and very limited design for specific properties has been achieved.In this current work, the uniaxial compressive properties of a 3D re-entrant lattice structure were modeled. Compressive Ti6Al4V and pure copper samples were manufactured via the electron beam melting (EBM) process and were then evaluated via mechanical testing. The results support the validity of the analytical models, and the difference in behavior between the Ti6Al4V and the pure copper are also discussed.Modeling Of The StructureFig. 1a shows a 2D re-entrant lattice structure which can be readily extended into a 3Dstructure as shown in Fig. 1b. The design parameters for the structures are shown in Fig.

1a shows a 2D re-entrant lattice structure which can be readily extended into a 3Dstructure as shown in Fig. 1b. The design parameters for the structures are shown in Fig. 1c, which include the length of the vertical (H) and re-entrant strut (L), the re-entrant angle (θ), and the thickness of the strut. Fig. 1b shows a simple pattern where the struts have rectangular cross sections. In practice, the shape of the cross section may vary.a. 2D unit cellb. 3D unit cell in a pattern Fig. 1 Re-entrant lattice structurec. Design parametersFrom Fig. 1b it can be seen that the 3D re-entrant lattice structure is orthotropic, with two geometrically identical directions (x1 and x2) and a third direction (y) which is different than the other two. The auxetic behavior can be tailored to a greater extent in the y-direction due to the anisotropy of the structure.

The auxetic behavior can be tailored to a greater extent in the y-direction due to the anisotropy of the structure. The current work therefore focuses on mechanical properties in the y- direction.Consider a remote compressive stress (σ) applied in the y-axis direction on an infinite 3Dre-entrant lattice structure. Due to the structure’s symmetry, the relationship between the stress and the resulting force on the struts is simplified into the case shown in Fig. 2a, where the force components can be shown to be:F=1 2σ2(Lsin) θ2=2 L 2 σsin2 θa. Simplified structureb. Decomposition of the structureFig. 2 Analysis of the structure under compressive loading in y-axis directionFrom Fig. 2(b), the deformation of the structure is influenced by the bending of the re-entrant (e.g. diagonal) struts as well as axial compression of the vertical struts.

2(b), the deformation of the structure is influenced by the bending of the re-entrant (e.g. diagonal) struts as well as axial compression of the vertical struts. Under the small deflection assumption [16], it is determined that the re-entrant strut is subject to critical loading. For the simple case involving struts a square cross section of thickness t, the Poisson's ratio (vy), the structural modulus (Ey) and the yield strength (σy) of the lattice can be obtained using standard beam theory as follows:ν y−=2 L cos 2 2 t α( αθ 2 L +− sincos 2) θ θ(1)(2)2 ( σ Y−cos ) 2( − θ α σ = 2 2 2 L t sin + ε αE=yθ4y θ )229Fsin 4 64 t 4 σ Y3t22Fcos 9 F−2 σ Y(64−⎛ ⎜ ⎝L2t sinθθσ Y 2 sin 4 t2θ ) t⎞ ⎟ ⎠2=EFL 8sinθwhere α=H/L, and E and σ are the modulus and yield strength of the bulk material, respectively. Eq. (4) is an implicit equation and can be solved numerically. Eq.

Eq. (4) is an implicit equation and can be solved numerically. Eq. (4) indicates the critical compressive force for yielding, which can be converted into the yield strength as follows:σ=2LFm 2 sin2θSample design and preparationTwo different configurations were designed using the structural parameters shown inTable.1. The cross section of the struts were made square in the CAD model with different sizes for the vertical and re-entrant struts t1 and t2 respectively. The relative density (RD) was taken to be the mass of the lattice structure divided by the mass of solid block having the same bounding box dimensions. The theoretical Poisson’s Ratio (PR) values were calculated by Eq. (2).

The theoretical Poisson’s Ratio (PR) values were calculated by Eq. (2). As indicated in Table 1, design configuration D2 should exhibit more significant auxetic behavior due to its more negative PR value.Table 1 Design parameters for experimental samplesDesign H (mm) L (mm)D1 D27.6 124 6t1 (mm) 1 1t2 (mm)0.94 0.71θ (°)70 45RD0.1160 0.0915PR0.463 -1.496Six samples for each design configuration were produced via the electron beam melting(EBM) process using both Ti6Al4V and pure copper. Ti6Al4V samples were produced in an Arcam A-2 machine using -100/+325 mesh spherical Ti6Al4V plasma rotating electrode (PREP) powder. The copper samples were produced in an Arcam S-12 machine using -100 mesh gas atomized copper powder. In this study, the density (ρ), yield strength (S) and modulus (E) values for the solid Ti6Al4V and copper are taken as: ρTi6Al4V=4.43g/cm3, STi6Al4V=1050MPa, ETi6Al4V=114GPa, ρCu=8.94g/cm3, SCu=330MPa, ECu=119GPa.(3)(4)(5)a.

CAD modelsb. Samples produced by EBMFig.3 Design and manufacturing of samplesAfter parts were produced by the EBM process, the lattice block dimensions were measured using calipers, and the strut size was measured via optical microscopy. Sample weights were recorded using a digital scale. The recorded dimensions and weights with sample standard deviations are shown in Table 2. Two strut thickness values are provided for both Ti6Al4V and copper samples, indicating that the fabricated struts had rectangular rather than square cross sections. The dimensional variation for the strut size is more significant for copper than for Ti6Al4V. It is hypothesized that this was due to the greater degree of particle sintering at the part surface introduced by high thermal conductivity.

It is hypothesized that this was due to the greater degree of particle sintering at the part surface introduced by high thermal conductivity. This also leads to an increased relative density of the structure, which according to the theory of cellular solids, playing a significant role in the mechanical performance of the structures [17].The manufacturing quality of EBM for the Ti6Al4V parts is relatively consistent.However, due to the heat dissipation, some of the particles around the melted struts become lightly sintered to the struts. The surface quality of the EBM fabricated struts is therefore rather rough. Fig. 4 illustrates the deviation of strut size/form from the ideal CAD model representation introduced by the surface texture. In order to evaluate the mechanical properties, the minimum strut dimension d was adopted as the thickness values of the struts.

In order to evaluate the mechanical properties, the minimum strut dimension d was adopted as the thickness values of the struts. Furthermore, the cross section of the actual struts is apparently not an idealized square and is therefore another source of deviation between the experiments and the theoretical predictions.The process theme of the EBM process with pure copper has recently been developed inhouse with the implementation of the new multi-beam control, therefore copper was also used for this current study. It is worth noting that the process parameters has not been optimized for mesh structures. Due to its high thermal conductivity, it was expected that the dimensions of the copper samples would exhibit relatively large variation.Fig.

Due to its high thermal conductivity, it was expected that the dimensions of the copper samples would exhibit relatively large variation.Fig. 4 Surface of EBM strutsTable 2 Unit cell dimensions of the actual parts made by EBMDesignH (mm)L (mm)t1 (mm)t2(mm)θ (°)RDTi6Al4VD1D27.436±0.031 4.101±0.00811.945±0.016 6.238±0.0051.085±0.0441.076±0.0451.180±0.028 (1.078±0.076) 1.155±0.034 (0.815±0.020)70450.142±0.0050.115±0.002D1D27.255±0.170 4.197±0.02612.526±0.118 6.375±0.043Copper 1.209±0.146 (1.698±0.059) 1.507±0.087 (1.160±0.101)1.973±0.214 (1.106±0.150) 1.517±0.181 (1.045±0.115)70450.208±0.0050.155±0.002Experiments and resultsCompression tests were performed on an Applied Test Systems 1620C Universal TestingMachine, with maximum loading of 89000N. Two case hardened steel platens were used to perform the compressive tests.

Two case hardened steel platens were used to perform the compressive tests. An extensometer was attached directly to the platens, as they cannot be attached to the lattice block samples. Fig. 5 shows the layout of the experiments.a. Ti6Al4Vb. Pure copperFig. 5 Compressive tests of the samplesTypical compression strain-stress curves for Ti6Al4V and copper samples are shown in Fig. 6. In comparing the two materials, it is apparent that the Ti6Al4V samples exhibit a failure mode very similar to that of a brittle material. After the initial linear stress-strain region, the stress drops abruptly to a rather low value without apparent yield. During compression testing, the Ti6Al4V samples typically suffered catastrophic failure of one whole layer normal to the loading direction, which corresponded to the steep drop of stress value in the curve.

After the first failure, the stress level of the Ti6Al4V samples stays relatively low and stable with increasing strain before entering another linear stress-stress cycle. This corresponds to the loading and subsequent failure of another complete layer.As expected, strain-stress curves for the copper samples showed considerable ductility.After the initial linear stress increase, there exists an apparent yield period. Sample failure is characterized by gradual deformation of the structure without abrupt fracture of individual struts.a. Ti6Al4V - D1b. Ti6Al4V - D2c. Copper - D1d. Copper - D2Fig. 6 Representative stress-strain curves for the compressive samplesThe total energy absorbed during the compression was calculated for each sample. Theenergy absorption by each structure up to 25% strain was obtained by calculating the total areas under the strain-stress curves. This accounts for the total energy absorption of the first layer of the structure.

This accounts for the total energy absorption of the first layer of the structure. It was expected that ductile materials (e.g. copper) would show a significantly improved energy absorption ability compared to the relatively brittle Ti6Al4V samples. This is reflected by the curves in Fig. 6.The mechanical properties as well as energy absorption characteristics for the Ti6Al4V and copper samples are shown in Table 3.

This is reflected by the curves in Fig. 6.The mechanical properties as well as energy absorption characteristics for the Ti6Al4V and copper samples are shown in Table 3. Significant variation of values was observed for the D2 samples made by copper, which was likely caused by the greater degree of geometrical irregularity at the ends of the struts.Table 3 Mechanical properties of the samples under compressionDesignStrength (MPa)Modulus (GPa)Energy absorption per unit volume (MJ/m3)Energy absorption per unit weight (kJ/kg)Ti6Al4VD1 D231.765±1.309 12.898±0.5082.231±0.567 0.646±0.0820.622±0.234 0.432±0.0420.984±0.343 0.851±0.092CopperD1 D22.804±0.513 0.624±0.2040.117±0.054 0.021±0.0080.188±0.017 0.057±0.0100.101±0.011 0.042±0.007DiscussionBased on the measurement results of the actual samples, the theoretical prediction of thecompressive strength as well as the modulus was obtained.

Because of the variation of the design parameters, there exists a range for the resulting mechanical properties, which is shown in Fig. 7, along with the comparison between the theoretical prediction and the experiments. It could be seen that the prediction for Ti6Al4V samples is in relatively good agreement with the experimental results, while the prediction for copper samples is significantly higher than the actual values.a. Strength-Ti6Al4Vb. Modulus-Ti6Al4Vc. Strength-Copperd. Modulus-CopperFig. 7 Comparison between the model and the experimentFor Ti6Al4V samples, the theoretical prediction shows the same trend as the actualexperimental values, with the deviation between the prediction and the actual values relatively consistent. This deviation could be largely attributed to the reduced effective length of the re- entrant struts. As shown in Fig. 8 the thickness of the struts causes the actual length of the re- entrant struts that is deflectable to be shorter.

As shown in Fig. 8 the thickness of the struts causes the actual length of the re- entrant struts that is deflectable to be shorter. This effect becomes more significant as the re- entrant angle becomes smaller. It is difficult to determine the effective length accurately. For the current work, however, it could be roughly estimated that the reduction of the length was between 0.7-0.9mm based on the design configuration. After the correction was applied, the theoretical predictions show considerable improvement of match with the actual values, as also shown in Fig. 7 (a)-(b).Fig. 8 Demonstration of the reduced effective length for re-entrant strutsFor the copper samples, the large difference between the theoretical prediction and theactual results is believed to be largely attributed to large deviations in strut dimensions. From the measurements shown in Table 2, the copper strut dimensions vary by more than 13%. From Eq. (3) and Eq.

From the measurements shown in Table 2, the copper strut dimensions vary by more than 13%. From Eq. (3) and Eq. (4), the strut dimension (t) has a significant effect on the mechanical properties. Fluctuations in strut dimensions will therefore result in large variations in mechanical properties as shown in the error bars in Fig. 7 (c)-(d). In addition, significant layering effect could be observed for the copper parts, as shown in Fig. 9. This indicates potentially weak inter-layer metallurgical bonding. Also visible in Fig. 9 is a large void defect which is marked by the arrow. From the strain-stress curves of some copper samples, it could be reasonably concluded that the yield of some of the struts occurs even at the very early stages of compression, resulting in the curving of the strain-stress curve. With this explanation, it could be estimated that the actual effective strut dimensions for D1 and D2 copper samples are around 0.5-0.7mm.

With this explanation, it could be estimated that the actual effective strut dimensions for D1 and D2 copper samples are around 0.5-0.7mm. The manufacturing defects could in turn be explained by the large coefficient of thermal conductivity of copper. Due to the severe dissipation of thermal energy during the melting process, it is more difficult to control both the metallurgical integrity and the dimensional accuracy. In the case of current study, both were compromised to a large extent.Fig.9 Strut of the copper samplesFor most stochastic cellular structures, the strain-stress curve shows an apparent plateaustage after the onset of initial yield [17], which was not observed in all the Ti6Al4V samples, and most of the copper samples.

It is known that most of the energy absorption of the cellular structures occurs at the plateau stage, therefore, it could be reasonably expected that the energy absorption ability of the re-entrant lattice samples made by Ti6Al4V was significantly limited. In comparison, copper has much higher ductility, and its strain-stress curve also shows much smoother pattern. Comparing the energy absorption of the Ti6Al4V samples shown in Table 3, design configuration D2 exhibits much lower strength and modulus while absorbing comparable energy per unit weight compared with D1. It is preferred for the structure to exhibit lower peak resistance force in many energy absorption applications, therefore, in the applications where the total weight is limited, designs with greater auxetic behavior (e.g.

D2) will be favored.After the failure of two layers, the copper samples started to densify, which could becharacterized by the steady and mostly linear increase of the stress at the end of the strain-stress curves shown in Fig. 6 (c)-(d). Due to the copper’s greater ductility, individual struts did not always fracture and detach from the structure. They therefore act as resisting sites for further deformation and facilitate the onset of densification. Fig. 10 shows one of the D2 copper samples after compression. It could be seen that the struts have formed a stable frame structure that behaves more like a densified solid that it would had the struts fractured cleanly from one another.Fig. 10 Compressed copper sampleConclusionIn this work, a 3D model for a re-entrant lattice structure was presented in which thelattice geometry was defined as a function of four parameters.

10 Compressed copper sampleConclusionIn this work, a 3D model for a re-entrant lattice structure was presented in which thelattice geometry was defined as a function of four parameters. Two different design configurations were created for comparison in compressive tests using relatively brittle Ti6Al4V and ductile pure copper. Samples were manufactured via the EBM process and evaluated through physical tests. The results were compared with the theoretical predictions, and corrections of the theory were made to compensate for deviations introduced by the manufacturing process.The results for the Ti6Al4V samples are in reasonably good agreement with the analytical models. On the other hand, the results for the copper samples show consistently large deviation from the predictions, which could be largely attributed to the small effective strut size caused by manufacturing variations such as voids and the step effect.

The large thermal conductivity of the copper results in significant sintering effect which contributes to the diminished surface quality of copper parts in the EBM process. More process optimization is therefore needed with this material to improve surface quality and accuracy.With greater auxetic behavior, the energy absorption per unit weight increases while the maximum react strength (e.g. resistant force) could be controlled in a low level. This makes the structures with large auxetic behavior favorable in the energy absorption applications where weight is a limiting criteria.References[1] Roderic Lakes. Foam Structures with a Negative Poisson's Ratio. Science. 235(1987): 1038- 1040. [2] F. Scarpa, P. J. Tomlin. On the transverse shear modulus of negative Poisson's ratio lattice structures. Fatigue & Fracture of Engineering Materials & Structures. Vol 23, 2000, p 717-720. [3] F. Scarpa, G. Tomlinson.

sign and Construction of a 6-DoF Fabrication PlatformJames Grossman, Warren Parad, Hod Lipson Mechanical & Aerospace Engineering, Cornell University, Ithaca NY 14853, USAAbstractThis research demonstrates a working freeform fabricator with a six degree of freedom print- head capable of additive fabrication onto existing structures. A parallel actuation mechanism was developed with stationary motors. Kinematic simulations of the printer’s motion were used to analyze and optimize the design. The components of the printer are described, including the control system and print head mechanism. A working fabricator was then constructed and tested. An additive three dimensional structure is demonstrated and the accuracy and reliability of the printer is analyzed.Goals and motivation The goal of this project is to create a printer capable of additive construction onto existing parts or structures (Figure 1).

While there are many types of three dimensional printers that are capable of building a part from scratch layer by layer, there has been relatively little work done on printers with the adaptable print-heads and robust control mechanisms necessary to print onto an existing surface. Existing bodies of work tend to focus on regular planar deposition systems with a built-in CNC component allowing for out of plane removal (e.g. [12]), although there have been successful attempts at parallel mechanisms for subtractive assembly (e.g. [4,11]). There are many uses for such a device: In order to perform touch up or repair work, as well as for augmenting existing parts (e.g. [1]). A flexible build platform can help prevent damage to existing parts and achieve an optimal angle for deposition of material for better functional performance.

[1]). A flexible build platform can help prevent damage to existing parts and achieve an optimal angle for deposition of material for better functional performance. It is also advantageous to be able to print in situ, especially when the site is not normally accessible (e.g.[3]).Figure 1: Completed Printer: CAD model (left), actual printer (right)A secondary goal of this project is to investigate printer systems with parallel actuation. Parallel actuated systems are ones in which the motors that control the system are independent; these motors are fixed in place and do not move during operation of the device. They provide a better tolerance than serial actuators, since an error in one rail will not necessarily affect the other rails, whereas on a serial mechanism such errors accumulate.

They provide a better tolerance than serial actuators, since an error in one rail will not necessarily affect the other rails, whereas on a serial mechanism such errors accumulate. Parallel mechanisms are known for their rigidity but sometimes suffer from severe nonlinearities and singularities.The fundamental architecture of additive manufacturing technology used today has not evolved substantially since their inception. The majority of rapid prototyping platforms use a three degree of freedom print head that follows the Cartesian coordinate system. These printers are capable of moving in the x, y, and z directions, but are unable to control the pitch, yaw, or rotation of the print head. A six degree of freedom print head would provide an increased level of control, making it very useful for in-situ printing, e.g.

A six degree of freedom print head would provide an increased level of control, making it very useful for in-situ printing, e.g. work on biological systems where the substrate extruded by the print head needs to be placed in exactly the right position and orientation.Method/ Technical Approach There is a significant body of work on analysis and control of parallel manipulators. The majority of technical papers reviewed detail the analysis of analogous systems such as a Stewart platform (e.g. [2,5,10,9,6]). These methods are based on analysis the forward kinematics of the platforms, which describe the position and orientation of the platform given the lengths of the actuators. This is considerably simpler than inverse kinematics, which requires the calculation of the actuator lengths based on the desired position of the platform.We used a relaxation-based approach to perform the inverse kinematics [7].

Starting with a given target print-head position and orientation, we mathematically relax the actuator positions until they reach in equilibrium. This relaxation process is then repeated for multiple points along a desired path of the print head, yielding the required trajectory of the actuators for that path.A critical challenge in the design of a parallel actuator is its sensitivity. Because of the nonlinear transfer function between actuators and the end effectors, some regions may be very sensitive to actuator perturbations (high gain), while other regions may be insensitive to actuator positions (uncontrollable), required discontinuities, were entirely unreachable (dead zones), or were undefined (singularities). Using the relaxation process, we could determine, for each point along a path, the sensitivity of the print-head’s position to perturbations in the actuator positions.

Using the relaxation process, we could determine, for each point along a path, the sensitivity of the print-head’s position to perturbations in the actuator positions. This calculation yielded a sensitivity score that could serve as a figure of merit for the quality of a given design (configuration of motor positions and rail orientations). To this end a series of scripts were written that could be used to easily tweak the basic design parameters of the printer, such as the length of the rails or the shape of the base, in order to determine which combination of shapes and lengths results in the most stable design. Stability in this case is defined as the ability to duplicate a given design within a margin of error and a reasonable time frame.

Stability in this case is defined as the ability to duplicate a given design within a margin of error and a reasonable time frame. To achieve stability the final printer design must minimize any internal singularities and be geared towards producing a wide variety of basic geometric paths and shapes in the largest build space.Once the basic design parameters have been optimized, we developed a three dimensional model of the printer that represented the layout and dimensions of the components. As construction progressed the original model was updated to take into account minor changes that were made. Several of the drive components, as well as the print head and rails were based on the Fab@Home system [8].Testing of the printer involved execution of paths using the printer of increasing complexity in order to demonstrate the full capabilities of the device.

Dry printing runs of basic three dimensional shapes were attempted and the results were documented.Implementation and Results We used a bruit-force search to explore the space of parameters relating to the shape of the base, middle, and top of the printer as well as the length of the rails and arms and the overall diameter of the various levels of the printer. Using these dimensions as parameters the program attempted to draw a predetermined shape by relaxing the lengths of the rails.After each run, the program provided the values of the various design parameters changed during each iteration, the amount of time that it took to produce the completed target path and a number relating to the accuracy of the final path compared to the desired path.

Since the design space was so large owing to the number of components that could be modified, we manually identified trends in order to avoid groups of possible designs that were exhibiting large number of internal singularities or other adverse features.Figure 2 shows the kinematic model and resulting convergence curves for a search process for a configuration where the leg changes needed to create a design did not require any instantaneous shifts. Another script was written that calculated the lengths of the arms required to reach every point in the design; this script also tested the provided design to make sure that none of the points exceed the arm length specified.

Unfortunately, the accuracy required for this project meant that the synchronous version of the relaxation algorithm was needed; this proved to be the largest performance bottleneck on the tests, though the effects were minimized by dividing the calculations over several computers.(a)(b)Figure 2: Design optimization process and final optimized design: (a) kinematic model (b) convergence of rail lengths(a)(b)(c)(d)Figure 3: Construction details: (a) the print-head (green) connected to the six rails using ball-and- socket links; (b) top of the rail; (c) Butterfly Tabs: These butterfly tabs keep the two base pieces aligned and the thermoplastic inserts with feet keep the printer level; (d) Bottom of the railThe next step in this project was designing a CAD model of the printer. We used 6mm acrylic sheets and 3D printed ABS plastic parts to form the frame of the printer, with connections and offsets comprised of various steel and brass components.

We used 6mm acrylic sheets and 3D printed ABS plastic parts to form the frame of the printer, with connections and offsets comprised of various steel and brass components. The design model was divided into six sections: the base, the rail system, the travels for the rail, the top connecting plate between the rails, the electronics board and the print-head itself (adopted from Fab@Home).The base section was comprised of two acrylic sheets joined together with a butterfly-tab and acrylic locking system (Figure 3c). Once the parts were cut out, thermoplastic inserts were placed into the base so that rubber feet could be screwed onto the bottom. These feet allow the base to be adjusted so that it is level as well as protecting the acrylic from the surface that the printer is resting on.

These feet allow the base to be adjusted so that it is level as well as protecting the acrylic from the surface that the printer is resting on. Additionally, the rubber feet act as vibration absorbers, preventing minor vibrations on the table from affecting the printer and vice-versa.Directly onto top of the base is the rail system for the printer. The rail comprised six identical ACME threaded rods connected to stepper motors and mounted on ABS plastic frames. The frames orient the motors in the correct direction while also providing a base for the support rods, which help guide the travel sections up and down the rails. Figure 3 shows the rail system (Figure 3a), the bottom anchor with motor (Figure 3d) and the top anchor (Figure 3b). One of the main issues that arose with these parts was the tolerances of the printed ABS plastic and the connections between it and the stepper motors.

One of the main issues that arose with these parts was the tolerances of the printed ABS plastic and the connections between it and the stepper motors. This issue was eliminated by drilling out theexisting mounting holes in the ABS plastic so that the bolts holding the motor could be connected without straining or deforming the part. The precision steel support rods were attached to this frame using a combination of mounting holes and shaft collars that sit on either side of the plastic mount, preventing the rods from sliding (see Figure 3d). The final component of the rail system is the attachments for the limit switches on the top of the mounting frame (shown in Figure 3b). These switches are connected to the electronics board and prevent the stepper motor from pulling the travel too close to the base.The travel is the device that connects the rails of the printer to the print-head itself (bottom-left of Figure 3a).

The design of the travel was one of the first parts finished for this printer, and the limitations on its size helped to determine the dimensions for the rest of the parts. The travel is constructed from acrylic sheets held together with a combination of press-fit joints and the tab-notch system used on the Fab@Home printer. Additionally, the travel includes oil impregnated brass bearings and a plastic threaded nut and flange assembly which connect to the support shafts and the ACME rod respectively. The connection to the print-head is done by a custom built ball and socket joint that is attached to the front of the travel. This joint is made from rapid-prototyped ABS plastic socket into which a 20 mm bead is press fit (see red beads in Figure 3a).

This joint is made from rapid-prototyped ABS plastic socket into which a 20 mm bead is press fit (see red beads in Figure 3a). This bead has a thermoplastic insert melted into it, allowing a threaded rod to be attached between the sockets on the print-head platform and the ones of the travel, forming the ‘arms’ of the printer.To prevent the rail system from vibrating during operation the printer also incorporates a top plate that connects the rails together allowing them to provide structural support for one another (seen in Figure 3b). This top plate was not part of the original model of the printer, and it proved to be the most difficult component of the printer to design due to the irregular angles between the top of the rails. The plate is comprised of three ABS plastic parts that bridge the gap between the adjacent rails and an acrylic sheet that connects these components to one another.

The plate is comprised of three ABS plastic parts that bridge the gap between the adjacent rails and an acrylic sheet that connects these components to one another. The bridging parts also have low friction ball bearings on them to reduce the friction between the ACME rods and the ABS plastic, as well as shaft collars around the rail’s steel support rods to prevent the top plate assembly from slipping. Finally, the top plate contains mounting brackets for limit switches to prevent the travel from impacting the top of the printer.The control system for this printer is similar in form to the system for the Fab@Home Model 1; it is comprised of stepper motor amplifier boards connected to a microcontroller board with header connections and a USB port, allowing the system to be controlled by any modern computer.

These components are mounted onto an acrylic board with standoffs to allow for air flow around the parts, and the whole board is mounted vertically on one side of the printer’s base. To ensure that the board does not fall over, support struts made from acrylic provide additional connections between the base of the printer and the back of the electronics board.The final component of the printer is the print-head itself (Figure 3a). This part is comprised of a slightly-modified syringe tool mount from the Fab@Home system which is attached to a custom acrylic base. Similar to the travel section, this base provides the connections between the print-head and the arms of the printer which are in turn attached to the travel and rail systems. In order to get the best range of motion out of the ball and socket joints on this base there are acrylic offset pieces that hold the ball joints at specific orientations to avoid lockups in the arms or collisions between two arms.

Ideally the print-head should be as light as possible to reduce the strain on the rails system. As such one goal when designing this part was to reduce the weight as much as possible. The main weight comes from the stepper motor used to extrudematerial through the syringe tool; however effort was made to cut out as much acrylic as possible from the attachment plate.After the printer was assembled, tests were performed in order to ensure that the physical system matched the simulated results. These tests included executing free form paths with and without a syringe attachment, demonstrating the capability to move in a controlled manner through all six degrees of freedom and small print tests in which the ability to deposit onto an inverted sphere was demonstrated. Next, tests were conducted to determine the accuracy of the system. The first test involved printing straight lines on wax paper.

Next, tests were conducted to determine the accuracy of the system. The first test involved printing straight lines on wax paper. The linearity of these lines was then determined using a first order polyfit function in MATLAB using the center points of the lines as the experimental data. While straight lines are trivial for a gantry system, they are as difficult for a parallel system as any other curve. Linearity therefore serves as a good test of system precision.(a)(b)(c)(d)(e)(h) Figure 4: Straight line test results.

Linearity therefore serves as a good test of system precision.(a)(b)(c)(d)(e)(h) Figure 4: Straight line test results. Images were cropped and rotated; best fit lines were calculated and placed by MATLAB.(f)(g)(i)(j)Line Variation (mm) R2 value(a) 0.37(b) 0.25(c) 0.24(d) 0.31(e) 0.41(f) 0.55(g) 0.83(h) 0.82(i) 0.170.9518 0.9964 0.9538 0.5630 0.6083 0.7268 0.9693 0.9522 0.9986 0.9980 Table 1: Average variation from mean in mm and R2 values of best fit lines.The accuracy of the lines was measured by calculating the average variation from the mean, and the r-squared values of the fit.

The results are summarized in Table 1; with an average deviation of 0.42±0.075 mm and a mean R2 value of 0.87±0.05.(j) 0.25Figure 5: System test: (a) Complete structure test: A silicone-based caulking material was used to form a cube on top of an existing hemisphere; (b) Reliability Test: The hash marks are 1/64” apart.The tests were performed using a syringe deposition tool filled with a silicon based tile sealant. The varying thickness of the lines is caused by the inaccuracies in the manual control of the syringe plunger. Target length for the lines was 1.5 inches – this was fairly close to the actual length; the images above show that the lines generally fit in this dimension with only a little creep due to the syringe tool at either end of the line.The printer’s ability to fabricate a three dimensional structure was tested by printing an open ended cube onto a hemisphere formed from half of a ping-pong ball (Figure 5a).

The test was performed by manually instructing the printer to follow a series of parametric coordinates that defined a rising square pattern, while taking into account the curvature of the hemisphere. This allowed the printer to form the cube in a single print operation.A repeatability test was also performed to determine the printer’s ability to place dots on a sheet of paper in a repeatable fashion. The test involved placing two dots one inch apart on a flat surface thirty times in a row. The dots were made by a pen mounted in the syringe tool attachment with the results shown in Figure 5b. The dot in the center of the image is a test dot demonstrating the size of the dots made in the run. As the image shows, the print-head demonstrated a regular precession during the test that prevented it from placing the dots exactly on top of one another.

As the image shows, the print-head demonstrated a regular precession during the test that prevented it from placing the dots exactly on top of one another. However, since the distance between the dots remained constant during this test the precession was likely due to the slight deflection to the print-head caused as the dots were placed. All these tests were done using open-loop control. We expect that higher precision could be attained using closed loop system.Conclusion The purpose of developing a six degree of freedom rapid prototyping platform is not just to expand the design space of rapid prototyping machines, but to create a printer that is capable of more than the current state of the art. The printer detailed above is capable of drawing arbitrary three dimensional shapes through the use of a rotating and translating print-head.

The printer detailed above is capable of drawing arbitrary three dimensional shapes through the use of a rotating and translating print-head. The use of a parallel actuation system to drive the printer demonstrates this platforms use in the free form fabrication field.Acknowledgements This project was made possible due to funding from the ELIA Life Technologies group and the Engineering Learning Initiatives grant from Cornell University.References1. Chung M., Malone E., Tolley M. T., Chepaitis A. J., Lipson H., (2008) “Object Augmentation for the Visually Impaired Using RP”, Proceedings of the 19th Solid Freeform Fabrication Symposium, Austin TX, Aug 20082. Dafoui, El-Mouloudi, (1998) “Analysis and Design of a Six-DOF Parallel Manipulator, Modeling, Singular Configurations, and Workspace”, IEEE Transactions on Robotics and Automation, Volume 14, No. 13. Grenouilleau J. C., O. Housseini, and F.

e Design and Construction of a Medical System to Optimize the Endoscopic Ultrasound ProcedureDarius P. Daruwala Christopher M. Johnson Donald R. Jacobsen*Rapid Prototyping Center, Milwaukee School of EngineeringMedical College of WisconsinABSTRACTThis project involved the use of rapid prototyping to produce a model of a section of the gastrointestinal (GI) tract which could be used for practice of the Endoscopic Ultrasound (EUS) procedure.Computed Tomography (CT) scans were obtained from Dr. Donald Jacobsen, Assistant Professor of Radiology at the Medical College of Wisconsin in Milwaukee. Apart from the final testing, the entire project was performed at the Milwaukee School of Engineering's Rapid Prototyping Center. To convert the CT scans into files that are compatible with the rapid prototyping machines, a software developed by Materialise, N.V., was used.

To convert the CT scans into files that are compatible with the rapid prototyping machines, a software developed by Materialise, N.V., was used. The rapid prototype models were used as master patterns for molds so that a polyurethane material with similar properties to human tissue could be used for actual simulation. Finally, these polyurethane models were placed in an enclosure and surrounded by a gelatin to simulate fatty abdominal tissue. The system was tested at Froedtert Memorial Lutheran Hospital under the supervision of Dr. Anthony Bohorfoush in conjunction with the Medical Physics Department of the Medical College of Wisconsin.INTRODUCTIONThe need for accurate models of the human body is greater in this age of rapid medical advancement. As new medical procedures and equipment are developed, few physicians and medical personnel have the knowledge and skills needed to use them effectively.

As new medical procedures and equipment are developed, few physicians and medical personnel have the knowledge and skills needed to use them effectively. Even the most skilled endoscopists have difficulty navigating the endoscope into the duodenum and targeting a specific organ. In addition, once the images are obtained, it is extremely difficult to interpret them due to the awkward orientation of the ultrasound transducer. Therefore, any system which will accurately simulate the human abdomen and allow physicians to practice EUS, will greatly increase the effectiveness of this relatively new medical procedure.The purpose of this project was to develop a system that accurately simulated the human abdominal organs when subjected to the endoscopic ultrasound (EUS) procedure.

Specifically, a model of the human stomach with the duodenum and an accurate simulation of the fatty tissue167that surrounds them were desired to allow physicians and radiologists to practice and teach the EUS procedure.DISCUSSIONRapid prototyping is a technique which is used to develop mechanical designs into a functional prototype within a few hours or days for physical or market testing. It can also be used to check the feasibility of new design concepts, to assess the fit of complex mechanisms, to make molds for wax cores in casting, and to use as a master pattern in silicon and epoxy molds.In medicine, rapid prototyping can be used in a process called reverse engineering, wherethe structure already exists but does not have regular geometric properties, such as that of the human body. Three-dimensional models can be created using two-dimensional data from imaging devices such as CT or MRI scanners.

Three-dimensional models can be created using two-dimensional data from imaging devices such as CT or MRI scanners. These three-dimensional models provide the surgeon with tactical visibility and an acute understanding in some complicated cases. A physical model of the patient's anatomy is easier to interpret than a series of two-dimensional images. The models can be used in surgical planning where complex operating procedures can be tried out on the model, greatly reducing the time in the operating room.

The models can be used in surgical planning where complex operating procedures can be tried out on the model, greatly reducing the time in the operating room. Physical models can also be used for communication of the patient's condition and expected outcome to members of the surgical team and for the patient's understanding of the treatment plan.ENDOSCOPIC ULTRASONOGRAPHYEndoscopic ultrasonography (EUS, sometimes simply called endoscopic ultrasound) is arelatively new technology that combines the features of two existing methods of obtaining structural information about the gastrointestinal tract and the tissues and organs immediately surrounding it.

The combination of endoscopy and ultrasonography (ultrasound imaging) allows physicians to obtain information that is not available or is not as accurate by other non-invasive means.With conventional ultrasound imaging, computed tomography (CT), and magneticresonance imaging (MRI) already successfully in use, why was an invasive ultrasonography technique needed? CT, MRI, and transcutaneous ultrasonography all have limitations for resolving small structures within and surrounding the gastrointestinal tract. CT uses radiation and MRI is fairly expensive. Conventional ultrasonography may not provide adequate resolution of the structural details of the GI tract due to the low frequencies used to provide deep penetration.

Conventional ultrasonography may not provide adequate resolution of the structural details of the GI tract due to the low frequencies used to provide deep penetration. Since the EUS probe is actually placed in the GI tract, adjacent to the target area, higher frequencies can be used (little penetration required) resulting in better resolution of the tissues of the GI wall and of adjacent structures.Conventional abdominal ultrasonography, CT, and MRI will continue to be an importantpart of the diagnostic procedures for patients with various gastrointestinal problems. However for several specific applications, EUS has proven itself superior to the other imaging methods. Diagnostic imaging with EUS is particularly valuable for examining endocrine tumors of the168pancreas. EUS has also been used to evaluate other pancreatic tumors, pancreatitis, and to find gall stones lodged in the common bile duct.

EUS has also been used to evaluate other pancreatic tumors, pancreatitis, and to find gall stones lodged in the common bile duct. In addition to pancreas abnormalities, EUS has allowed physicians to find and stage early gastrointestinal cancers more accurately, often preventing unnecessary surgery (to obtain biopsies). On occasions, when surgery is necessary, EUS can help to better prepare surgeons and patients for surgery. Ultimately, detection of cancers that are small and at an early stage will increase the chances of cure.In the short number of years that the EUS procedure has been used, it has proven to be a valuable complement to conventional ultrasonography, CT, and MRI in the examination of the gastrointestinal tract and the surrounding organs.DESIGN SPECIFICATIONSThe specification for the system was to accurately simulate the behavior of humanabdominal organs during the EUS procedure.

This requirement applied to the stomach, duodenum, and the fatty tissue that surrounds the internal organs. They had to be of similar size and shape of human organs and possess ultrasonic properties exhibited by human tissue.The stomach and duodenum were required to be made of a soft, flexible material withnearly the same density as water. The human stomach is very flexible and is normally in a collapsed state, unless food or drink has just been consumed. It was also necessary that the stomach be watertight, as it is inflated with water during the EUS procedure. It was important for the stomach wall to have a uniform density in applications involving ultrasonic waves. Therefore, it was necessary to either utilize a molding technique that produced a casting as one complete part, or if cast in halves, develop a way to attach them without altering the density.

Therefore, it was necessary to either utilize a molding technique that produced a casting as one complete part, or if cast in halves, develop a way to attach them without altering the density. The material selected for the stomach and duodenum from over twenty possibilities was a flexible polyurethane called SKINFLEX from BJB Enterprises. With this material, it was necessary to use platinum-based silicone as the molding material, since SKINFLEX will not cure correctly against tin-based silicones. After much deliberation, it was determined that the most effective method of producing the stomach was to cast it in halves and then glue the halves with the casting material.In addition to the stomach and duodenum, it was also necessary to accurately simulate fatty tissue in the abdomen. The first requirement was that the material be extremely soft and flexible. Also, the tissue must accurately simulate the ultrasonic properties of human soft tissue.

The first requirement was that the material be extremely soft and flexible. Also, the tissue must accurately simulate the ultrasonic properties of human soft tissue. This again involved near water density, and material that simulated the attenuation effect of the cells and cell walls. It was determined that an organic gelatin could best provide the density and texture required. Also, mixing tiny graphite particles into the material would provide the ultrasonic attenuation due to cells. A gelatin-water mixture of approximately 7.5% gelatin to 100% distilled water provided the texture and hardness (firmness) closest to that of abdominal tissue.The enclosure needed for the project was not one of the primary design focuses of theproject. The only requirement of the enclosure was that it be approximately the size of an average human's abdomen and that it be able to hold the gelatin without deforming.

The only requirement of the enclosure was that it be approximately the size of an average human's abdomen and that it be able to hold the gelatin without deforming. A small169valve and tube were attached to the end of the duodenum to allow these organs to be filled with water and drained when needed. Four small plastic balls were placed in the gelatin to act as targets during ultrasound imaging. These balls were measured prior to placing them into the enclosure so they could be compared to measurements taken from the ultrasound images. Fishing line and washers were used to weight the balls down, and keep them in place until the gelatin solidified. Finally, a wire mesh screen was also placed in the enclosure to act as another target.MATERIALS SELECTIONThis project required the selection of two materials: (1) a material to simulate the stomachand the duodenum, and (2) a material to simulate the fatty tissue that surrounds the stomach and other internal abdominal organs.

It was necessary that both materials be flexible and posses similar ultrasonic properties to the actual human organs.The ultrasonic properties that were of primary importance were the attenuation and thespeed of sound in these materials. The speed of sound is usually related directly to the density of the material. Therefore, materials with densities close to that of water (approximately 1.0 g/cm3 ) were selected. Because the ultrasound transducer on the endoscope is placed in almost direct contact with the stomach wall, the attenuation and other ultrasonic properties of the stomach were important. The main requirement for the stomach was that it be of a uniform density. However for the fatty material that surrounds the stomach, acoustic impedence was of great importance.

The main requirement for the stomach was that it be of a uniform density. However for the fatty material that surrounds the stomach, acoustic impedence was of great importance. Therefore, in future models, very fine powdered graphite will be added to the gelatin to simulate the attenuation of human fatty tissue caused by cells,.The maj or materials that were examined for these applications were epoxy, gelatin, latex,polyurethane, and silicone. Latex and epoxy were dropped from consideration early, because they break down in water, and material suppliers were difficult to locate. Numerous companies were contacted, and a great deal of information was gathered to allow the best decision between the remaining materials to be made. The following table summarizes the most appropriate offerings of companies contacted:Table 1: Summary of Material Offerings of Companies Contacted.

The following table summarizes the most appropriate offerings of companies contacted:Table 1: Summary of Material Offerings of Companies Contacted. Hardness 45+ Shore A 5+ Shore A Gelatin-Like 35+ Shore A 70+ Shore D 14+ Shore A 20+ Shore A Gelatin-LikeMaterials Available Polyurethane Polyurethane Polyurethane Gel Polyurethane Polyurethane Silicone Polyurethane GelatinCompany Air Products BJB EnterprisesApplication Stomach Stomach Fatty Surrounding Stomach Stomach StomachIMold Stomach Fatty SurroundingCiba-Geigy Eager Plastics GE Silicones Hapco Vyse Gelatin170After analyzing the literature and samples received, polyurethane emerged as the bestmaterial for the stomach and duodenum. Because of the wide range of physical properties available in polyurethanes, it was the material of choice. Polyurethane has a hardness that varies from "Shore A" to "Shore D ". Also, polyurethane has a density close to the density of water.

Polyurethane has a hardness that varies from "Shore A" to "Shore D ". Also, polyurethane has a density close to the density of water. An extremely flexible polyurethane called SKINFLEX(BJB Ent.) was selected for the simulation of the stomach and the duodenum.Figure 1: Silicone Rubber Mold of Two Stomach HalvesIn addition to the polyurethane, gelatin was explored as a possible material to simulatethe surrounding fatty tissue. With gelatin from the Vyse Gelatin Company, several samples were made. This gelatin produced a material with suitable hardness and density, but it was found to decompose when it was exposed to air. After several days in an open container, the gelatin samples grew fungus and dried up, becoming very brittle.

After several days in an open container, the gelatin samples grew fungus and dried up, becoming very brittle. However, after contacting the material vendor, it became clear that the addition of a preservative would increase the life of the system.ASSEMBLY OF THE ENTIRE SYSTEMAfter successfully producing the stomach model with the duodenum, the components hadto be brought together to make a complete, usable system. The enclosure used was a simple plastic waste basket with a small valve fastened on the back side, near the bottom of the container. A small rubber tube was attached to the end of the duodenum and to the valve to allow the stomach and duodenum to be drained periodically. On the other end of the system, a larger rubber tube (1.0" I.D.) was attached. The stomach, duodenum and esophagus were positioned in the enclosure, and four small plastic balls were placed near the stomach and duodenum to act as targets during testing.

The stomach, duodenum and esophagus were positioned in the enclosure, and four small plastic balls were placed near the stomach and duodenum to act as targets during testing. A course wire mesh was also placed inside the171enclosure to act as an additional target. The system prior to the addition of gelatin is shown in Figure 2.Figure 2: System prior to addition of gelatin.Finally, the gelatin powder was mixed with six gallons of water at a 7.5% to 100% ratioand poured into the enclosure, around the internal organs. After several hours, the gelatin solidified.TESTING THE SYSTEMThe final step of the project involved testing of the system with Dr. Anthony Bohorfoushat Froedtert HospitaL The probe was inserted into the stomach and the first portion of the duodenum ultrasound and video images were obtained.The ultrasound images can be frozen, and printed at any time. Figure 3 shows animage from one of the initial tests of the stomach.

Figure 3 shows animage from one of the initial tests of the stomach. The main purpose of this initial test was to determine how well the material simulated the wall of a human stomach. The black half ellipse on the top of the image is the ultrasound transducer, itself. In this case, the transducer is inside of the stomach, sending waves out. The top bright white line is where the ultrasound waves hit the inside of the stomach wall, while the lower white line shows the interface between the outside of the wall and the water (gelatin was not used in the initial test). After172looking at the images, Dr. Bohorfoush admitted that this material looked very much like a human stomach wall.Figure 3: Ultrasound image of ball next to duodenum wall.Figure 3 also shows one of the small balls placed under the stomach.

The diameter of the ball and the distance from the transducer and the ball were measured and were within 5% of the actual values.Another portion of the procedure involves real-time visual images. The images are displayed on a color video monitor and can help the endoscopist navigate through the organs, as well as see any deformities on the inner walls of the gastrointestinal tract. Like the ultrasound images, the visual images can be printed at any time, and the entire session can be recorded for review at a later date. Figure 4, on the following page, shows visual images obtained from inside of the stomach. The yellow ball can be clearly seen floating in the gelatin.

Figure 4, on the following page, shows visual images obtained from inside of the stomach. The yellow ball can be clearly seen floating in the gelatin. The bright light of the scope makes it possible to see right through the stomach and gelatin.173Figure 4: Images of yellow ball and duodenum entrance from inside stomach.CONCLUSIONThe medical system, developed in this project to allow physicians to practice and teachthe endoscopic ultrasound procedure, was a great success. Rapid prototyping has proven to be a useful tool in surgical planning, as exemplified in this case. The doctors who tested the system were extremely pleased with the project, and commented numerous times on how similar the images obtained were to the images obtained from an actual patient. This system has exciting possibilities for training both new as well as experienced users in the performance of the procedure, and in the identification of abdominal anatomy as imaged with the EUS.

This system has exciting possibilities for training both new as well as experienced users in the performance of the procedure, and in the identification of abdominal anatomy as imaged with the EUS. The doctors, radiologists, and students at the medical colleges now have an invaluable tool at their disposal for practicing the EUS procedure.174BIBLIOGRAPHYAmerican Society of Gastrointestinal Endoscopy. Understanding Upper GI Endoscopy. Http://www.asge.comAmerican Society of Gastrointestinal Endoscopy. Endoscopic Ultrasonography, Update February 1994. Http://www.asge.comBaker, T.M. Integration of3-D Medical Imaging and Rapid Prototyping to Create Stereolithographic Models. Australian Physical & Engineering Sciences in Medicine, Vol. 16, No.2. 1993.Bohorfoush, Anthony G., M.D. Advances in Imaging ofBile Ducts and Pancreas. Grand Rounds at Froedtert Hospital. 1996.Bronzino, Joseph D. The Biomedical Engineering Handbook. eRC Press, Inc. 1995.Digestive Disease Center.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesign and Development of a Multi-Tool Additive Manufacturing SystemGrady W. Wagner, Lindsey B. Bass, Daniel A. Rau, Scott B. Ziv, Mitchell S. Wolf, David L. Wolf, Yun Bai, Viswanath Meenakshisundaram, and Christopher B. WilliamsDesign, Research, and Education for Additive Manufacturing Systems LaboratoryDepartment of Mechanical EngineeringVirginia TechAbstractAdditive manufacturing (AM) makes complex parts with a single class of material. Each AM technology encompasses specific techniques and requires diverse components to selectively form each layer, which has segregated AM research by respective technologies. Multimaterial AM exists, but it is the same class of material with the same deposition tool.

Multimaterial AM exists, but it is the same class of material with the same deposition tool. To fully benefit from AM, researchers must explore the combination of multiple AM modalities and materials such that a multifunctional part may be fabricated using strengths of multiple technologies. While the methods for fabricating each layer differ, all of the AM technologies share the fundamental layer-based approach. By recognizing this universal similarity coupled with the desire to make multifunctional parts, a single system has been created to combine five different AM modalities. In this paper, the authors discuss the design and development of a multi-tool AM system that includes binder jetting, material jetting, vat photopolymerization, paste extrusion, and filament extrusion. Examples of multifunctional, multimaterial parts fabricated by multiple AM processes in a single integrated process are demonstrated.1.

Examples of multifunctional, multimaterial parts fabricated by multiple AM processes in a single integrated process are demonstrated.1. Introduction1.1 Functional organization of AMAdditive Manufacturing (AM), also commonly known as 3D printing, is the process of fabricating parts layer-by-layer. While there are several different types of AM technologies, all share a common process flow of turning a computer aided design (CAD) model into a 3D object. As shown in Figure 1, the CAD model is first converted to a standard tessellation (STL) file, which creates sets of stored triangle vertices and normal vectors. Software slices the file into cross-sectional 2D layers and develops a tool path to manufacture these slices. The AM system then sequentially follows these toolpaths layer-by-layer to fabricate each part. Functionally, the only difference among all AM technologies is how each layer is fabricated [1].20051Figure 1.

Functionally, the only difference among all AM technologies is how each layer is fabricated [1].20051Figure 1. Process flow for AM technologies showing commonality between different modalities.Each additive technology creates parts in different ways because they use different classes of materials. Vat photopolymerization uses a UV light source to selectively cure regions of a liquid photopolymer resin vat. Binder jetting uses an inkjet printhead to jet binder into a powder bed of either a metal, ceramic, polymer, or plaster powder. Material jetting jets liquid photopolymer droplets onto a substrate and cures it with an overhead UV light source. Filament extrusion heats a thermoplastic beyond its glass transition temperature before selectively creating primitives.

Filament extrusion heats a thermoplastic beyond its glass transition temperature before selectively creating primitives. Paste extrusion forces viscous pastes through a nozzle with a mechanism such as applied high pressure compressed air.AM can be an advantageous alternative to traditional manufacturing methods due to its ability to create complex, multimaterial geometries efficiently and economically; however, one of the greatest challenges facing AM is the lack of printable multimaterial parts due to the fact that a single toolhead can only manufacture a single class of materials. With the ability to utilize multiple classes of materials in a single print, engineers can design for a part to serve multiple functions.

With the ability to utilize multiple classes of materials in a single print, engineers can design for a part to serve multiple functions. Each AM technology utilizes different fundamental physics to selectively create building blocks of material, which results in certain materials only being able to be printed with a certain AM technology [1].1.2 Multimaterial and multi-modality AM systemsRecognizing that the process flow for creating a part with a layer-based method isconsistent across all AM technologies, there exists an opportunity to develop an AM system featuring tools of multiple AM technologies within one machine. A multi-tool AM system would allow users to create multi-functional products by using multiple AM technologies in a single print to fabricate multi-material parts. For example, the system could fabricate “smart” products that feature integrated sensing and actuation by extruding conductive pastes into a part with a20062photopolymer exterior.

For example, the system could fabricate “smart” products that feature integrated sensing and actuation by extruding conductive pastes into a part with a20062photopolymer exterior. This machine would also be able to create functionally-gradient materials by placing ductile materials beside hard materials.There has been an increase in development of systems combining AM and traditionalmachining, which can be considered a hybrid manufacturing process, containing both additive and subtractive methods. Arnold and authors have designed a machine integrating material extrusion, milling, and turning in a modular approach [2], and Keating and Oxman combined extrusion, milling, and sculpting through the use of a six-axis robotic arm positioning the workpiece to fixed tool mounts [3]. Zmorph has fabricated a system that has a single tool head slot with attachment options of various extrusion nozzles, a laser for engraving, or a CNC tool.

Zmorph has fabricated a system that has a single tool head slot with attachment options of various extrusion nozzles, a laser for engraving, or a CNC tool. A tool can be manually changed out during a print and replaced with a new tool to continue fabrication with the new attachment [4]. Hyrel 3D has created a three-axis gantry with swappable tools within a common interface. The modular platform allows multiple extrusion heads that handle materials of different viscosities to simultaneously manufacture multiple parts. Additionally, extrusion nozzles may be interchanged with lasers and routers for engraving, cutting, or routing purposes [5]. nScrypt has developed a high-precision nozzle to handle pastes and filaments, and assembling multiple nozzles onto a gantry combined with a “pick-and-place” tool has been shown to be advantageous for applications such as bioprinting, antennae, and circuit components [6].

These resources were used in the design phase to better understand how machines have integrated diverse manufacturing technologies within a common work space.Multi-material AM has been typically accomplished within a single technology. ThePolyJet process, commercialized by Stratasys, was the first to adapt multi-material printing in material jetting by simultaneously inkjetting two materials in dropwise deposition patterns [7]. This strategy allowed for flexible polymers to be printed directly adjacent to rigid polymers expanding the Design for AM (DFAM) possibilities. Sitthi-Amorn and authors have expanded the material jetting capabilities by developing a machine-vision-assisted platform for multi- material AM that can support up to ten materials simultaneously.

Sitthi-Amorn and authors have expanded the material jetting capabilities by developing a machine-vision-assisted platform for multi- material AM that can support up to ten materials simultaneously. The system provides multi- material design and fabrication capabilities to the computer graphics community, and it creates the ability to fabricate complex materials including fiber optic bundles, fabrics, privacy screens, and LED lenses [8]. More recently, multi-material printing has been achieved with vat photopolymerization. One such example is the system developed by Inamdar and authors who used a rotating carousel to house four vats, which expands AM applications in colored printing, embedded electronics, and tissue engineering [9]. Multiple material capability can also be seen with material extrusion technologies.

Multiple material capability can also be seen with material extrusion technologies. Malone and Lipson created a multi-material extrusion fabrication system featuring extrusion of liquids, gels, and pastes as well as thermoplastics through an open-source extruder design [10]. Wang and Liu used a nozzle with varied pneumatic parameters to extrude both metal and nonmetal inks [11]. Espalin and authors combined two filament extrusion printers onto one platform to enable the production of novel, multi-material thermoplastic parts. Multiple extrusion tips provided the opportunity to vary layer thickness and width to fabricate parts attaining desired tailored properties such as color, mechanical strength, weight, and thermal conductivity [12].20073Multi-material AM has recently been expanded to integrate multiple technologies together in one build thus combining strengths from different technologies.

Lopes and co-authors have designed a machine to combine vat photopolymerization and extrusion, which was used to conduct research in printed integrated circuits [13]. A system that combines vat photopolymerization and extrusion has also been developed by the Diyouware Initiative in Spain. Instead of keeping all inherent functional components of each AM technology, the system was designed with an extrusion nozzle to deposit low-viscous resins and a laser diode to selectively UV cure the material [14]. Rize has developed a printer that combines extrusion and material jetting to fabricate end-use parts. An extrusion nozzle and a print head are mounted to the printer, and each can be activated in a voxel-wise deposition strategy to create functional thermoplastic and ink combinations [15]. Raza and authors have developed a system capable of depositing low-melting-point alloys, elastomers, and UV-curable resins.

Raza and authors have developed a system capable of depositing low-melting-point alloys, elastomers, and UV-curable resins. The printer features two different jetting nozzles and one paste extrusion nozzle, and a test part included UV cure adhesive, rubber, and solder all on one composite object [16]. While these approaches work towards the goal of multifunctional parts, there have been no previous systems designed combining more than two AM technologies as summarized in Table 1. This multi-tool AM system incorporating five AM technologies will expand the scope of multi-material applications by providing the opportunity to combine several classes of materials.Table 1. Summary of current systems that utilize multiple, different AM technologies. The system presented in this paper, shown in the bottommost row, is expanding the capabilities of multi-material printing across multiple technologies.

The system presented in this paper, shown in the bottommost row, is expanding the capabilities of multi-material printing across multiple technologies. For simplicity, extrusion of liquids, gels, and other inks and resins are combined within the “Material Extrusion: Paste” category.First Author or CompanyZmorph Hyrel 3D nScrypt Malone Lopes Diyouware Rize Raza WagnerMaterial Extrusion: Filament X X X XXXMaterial Extrusion: Paste X X X X X XX XVat Photopoly- merizationX XXMaterial JettingX X XBinder JettingX1.3 Goal and contextThe overarching goal was to design, build, and evaluate a single AM system that allowsusers to print with a combination of five additive technologies: binder jetting, material jetting, vat photopolymerization, filament extrusion, and paste extrusion. As a research printer for AM20084materials discovery, an open-architecture approach has been employed to provide users with full control to adjust printing parameters.

As a research printer for AM20084materials discovery, an open-architecture approach has been employed to provide users with full control to adjust printing parameters. A multi-tool AM system has been manufactured to incorporate all of the technologies in one toolhead assembly with an adaptable platform for layer fabrication. One main LabVIEW interface was created to access controls for each of the technologies. The authors followed a systematic design process, and the highlights will be discussed in this paper. Section 2 outlines the design goals and specifications that influenced design decisions. Section 3 presents the system overview and key components of the detailed design of the multi-tool AM system. Results and discussion are shown in Section 4, and a summary is provided in Section 5.2. Design Goals and Target SpecificationsAlthough the five technologies are present on the system, they are not necessarily allcompatible.

Design Goals and Target SpecificationsAlthough the five technologies are present on the system, they are not necessarily allcompatible. Figure 2 displays the compatibility chart developed during the conceptual design. The black lines between technologies indicate a compatible link, i.e. that the two technologies can be used together for successful fabrication. For example, binder jetting will be incompatible with many technologies. A powder bed for binder jetting and resin vat for vat photopolymerization cannot keep materials separate. Filament extrusion and material jetting would likely not work with binder jetting because the deposited material would have difficulty adhering to loose powder. Paste extrusion, on the other hand, is compatible with all of the other technologies.Figure 2.

Paste extrusion, on the other hand, is compatible with all of the other technologies.Figure 2. Compatibility chart for the five technologies encompassed in the multi-tool AM system.20095With material compatibility kept in mind, a list of requirements and specifications to encompass both the general and technology-specific performance needs of the system. While some requirements were more straightforward, such as the capability of printing in each technology and switching between them quickly, other requirements, such as providing reconfigurable build volumes to enable processing low and high volumes of raw material, arose from demonstrated weaknesses of other AM systems.

At a system level, the machine was required to:Encompass multiple additive technologies (binder jetting, material jetting, vat photopolymerization, paste extrusion and filament extrusion)Allow for fast reconfiguration among technologies (<180 s)  Be able to print both high and low volume parts (reconfigurable build volume from 50 mm x 50 mm x 50 mm to 200 mm x 200 mm x 125 mm)Be stable when printing (<2 m/s2 vibrations)In addition, each technology has specific, individual requirements because final partsdepend on the unique parameters and capabilities of each separate technology. These requirements include:Operate with variable parameters  Manufacture small features  Produce small layer thicknessesSystem level target specifications are summarized in Table 2. All of these requirementsand specifications influenced key design decisions and testing results which will be discussed in the subsequent sections.Table 2.

All of these requirementsand specifications influenced key design decisions and testing results which will be discussed in the subsequent sections.Table 2. Engineering characteristics and target specifications.Requirement/Engineering CharacteristicUnits Marginal Value Optimal ValuePrints using material jettingY/NYesYesPrints using binder jettingY/NYesYesPrints using vat photopolymerizationY/NYesYesPrints using paste extrusionY/NYesYesPrints using filament extrusionY/NYesYesTime to switch between technologiessec1805Build volume dimensionsmm50x50x50200x200x200Vibrations when printingm/s22.00.1201063. System Overview and Detailed Design3.1 System OverviewA detailed CAD design of the system and the fully-built system is presented in Figure 3. The main frame was manufactured out of 80/20 aluminum extrusions with diagonal braces at the corners for rigidity.

The main frame was manufactured out of 80/20 aluminum extrusions with diagonal braces at the corners for rigidity. Since the build volume will hold liquid resins, it was designed to be stationary to avoid sloshing while the toolhead moves along X and Y high load, high precision linear stages (Zaber A-LST-E series) that are bolted to cross beams of the main frame. The build plate is actuated in the Z direction to avoid axis collisions while maximizing build size and minimizing additional overall system size. The frame and build platform were bolted to a modified wooden table that allowed linear actuators to control Z motion of the build and feed chambers. A large sheet of high density polypropylene was bolted to the tabletop to serve as the main work surface, which would be compatible with any of the powders, resins, or other liquids used.

A large sheet of high density polypropylene was bolted to the tabletop to serve as the main work surface, which would be compatible with any of the powders, resins, or other liquids used. To accommodate wires and electronics, a shelf was attached to the back of the machine where most of the electronic boards could be protected from harmful materials. Wire looms and cable carriers were also added to reduce the risk of wire tangles with all of the moving components. The authors then decided that the toolhead and build platform operate independently, which allowed them to be treated as separate subsystems.Figure 3. (a) Detailed design of the multi-tool AM system in CAD. (b) Manufactured version of multi-tool AM system.3.2 Toolhead DesignA toolhead configuration had to be selected that allowed all five technologies to be incorporated, as well as fast reconfiguration between them. A secondary objective was a modular design that allows for integration of additional, new tools.

A secondary objective was a modular design that allows for integration of additional, new tools. To fulfill these requirements, a toolhead was designed featuring all individual technologies mounted on one rigid frame due to simplicity of machine integration, as seen in Figure 4. Each individual toolhead mounts to a lead screw actuated by a stepper motor which allows lowering into optimal printing position and retracting when finished, as shown in Figure 5, to avoid interference with printed parts. Future20117development can easily be implemented by creating a new bracket to the leadscrew if a new toolhead is to be included.Figure 4. CAD model of the toolhead with unique all-in-one design.Figure 5.

CAD model of the toolhead with unique all-in-one design.Figure 5. CAD models of (a) Filament Extrusion, (b) Paste Extrusion, (c) Material Jetting, and (d) Binder Jetting.Figure 5 provides detailed schematic of each individual toolhead:Filament extrusion (Figure 5a): AV6 hot-end by E3D allows printing of thermoplastic filaments at extrusion temperatures up to 400 ̊C. The robust and well-tested design includes a sharp thermal break to ensure filament will not be melted prior to entering the hot end. Paste extrusion (Figure 5b): A Nordson Ultimis V was selected because it offers precise starting, stopping, and control of applied and vacuum pressures.

- Material jetting (Figure5c): A Nordson PicoPulse jetting valve for materialjetting offers adjustable waveform and frequency, a jettable viscosity range of 1-2012810,000 cP, and high accuracy and precision.Binder jetting (Figure 5d): An HP C6602A thermal inkjet printhead has the ability to send controlled electric pulses to each individual nozzle.Vat photopolymerization (not shown in the figure): An LC4500UV projector provides high resolution projections for UV curing, and it is mounted to the rear of the toolhead.With all of the toolheads mounted to a single frame, the time to change between technologies is minimized.3.3 Platform DesignThe key build platform requirement was to be compatible with all intended materials,which includes, but is not limited to, powders for binder jetting and photopolymer resins for vat photopolymerization. To achieve this, a metal build volume was welded together with an actuator going through a seal in the bottom.

To achieve this, a metal build volume was welded together with an actuator going through a seal in the bottom. Metal allows compatibility with corrosive resins and the ability to add heating pads to the walls. Heating pads are advantageous to remove humidity in powders, cure jetted binders, reduce thermoplastic part warpage, and reduce viscosity of highly viscous resins. A feed chamber, collection chamber, side overflow chambers, and a recoat system were included for the binder jetting modality. The recoater uses a rotating roller to increase powder flowability and compaction and an overhead heater to cure binders. The recoat frame is adjustable to ensure spread layers of powder are smooth and level. These components integrated into a single platform design is displayed in Figure 6.Figure 6. CAD model of the platform design.A secondary requirement was to be reconfigurable for both research and production sized prints.

CAD model of the platform design.A secondary requirement was to be reconfigurable for both research and production sized prints. The build volumes (8”x 8” x 5” in Figure 6) are removable allowing the user to insert any size build volume into the machine. The swappable strategy where the build volume inserts over the actuator followed by the build plate is shown in Figure 7.20139Figure 7. Swappable build volumes: 5” x 5” x 5” shown.In addition to swappable build volumes, the build plates also had to be designed toaccount for different technologies. For example, extrusion requires a heated build plate for the best adhesion of filament to the surface, material jetting requires an anodized surface to avoid light reflection, and binder jetting requires a sealed build plate for the containment of powder.

To account for this, multiple build plates were manufactured, and a tight-tolerance clevis bracket and pin connect an attachment plate to the linear actuator that controls Z axis movement. The actual build plates bolt to tapped holes in the attachment plate with spacers in between to provide room for the silicone heating pad needed for extrusion as seen in Figure 8. The bolts are countersunk to ensure there is no protrusion on the build plate that could cause a fatal crash of the toolhead.Figure 8. CAD model of the build plate attachment. Filament extrusion build plate is shown.3.4 Control SystemWith such a complex system, there are many different aspects of the machine that need tobe controlled such as toolhead actuation, nozzle firing, and temperature regulation. To accomplish this, the authors decided to use separate microcontrollers (Arduinos) synced to a central control system. Each microcontroller is designated to specific tasks as shown in Figure 9.

To accomplish this, the authors decided to use separate microcontrollers (Arduinos) synced to a central control system. Each microcontroller is designated to specific tasks as shown in Figure 9. Since temperature regulation and controlling Z position require constant monitoring, one microcontroller constantly accomplishes this without disrupting another system. The microcontrollers communicate with LabVIEW subVIs controlled by a master VI. Additionally, LabVIEW directly controls the stages and projector.201410Figure 9. Control system hierarchy for all subsystem controllers and respective tasks.The Master VI combines all of the subVIs for each system into a single interface for all processes as seen in Figure 10. All of the process-specific settings can be set, such as number of passes on a binder jetting head or delay between projected images. COM ports for all of the microcontrollers and stages, temperature settings, and motion can also be controlled from this screen.

COM ports for all of the microcontrollers and stages, temperature settings, and motion can also be controlled from this screen. Within this system, the Master VI sends Serial Data to both of the linear stages for motion, as well as data, start, and stop commands to the Arduinos that control the subsystems.Figure 10. The LabVIEW front panel for the Master VI.2015114. Results and DiscussionWith the machine fully operational, general and technology-specific requirements andtarget specifications were tested to evaluate system performance.4.1 System-level evaluationThe system level requirements were tested separately from the individual technologyrequirements. Binder jetting,vat photopolymerization, filament extrusion, and paste extrusion are all successfully designed for and included in the machine. Material jetting has been designed for, but has not yet been incorporated.

Material jetting has been designed for, but has not yet been incorporated. The ability to print both low and high volume parts has also been successfully included with reconfigurable build volumes. In addition, reconfigurable build plates allow for all technologies to print in the same build volume, minimizing overall system size.In order to test the fast reconfiguration among technologies requirement, a stopwatch wasused to measure the time from when one toolhead finished depositing material until the next began laying material. The two toolheads used were filament and paste extrusion for maximum time since both toolheads need actuated, as opposed to other technology combinations such as paste extrusion and vat photopoymerization because the projector does not need actuated in the Z direction.

Results showed the time to retract one toolhead and lower a separate toolhead into printing position was 18 s.To ensure the machine was stable when printing the authors placed an accelerometer onthe frame. To simulate worst case conditions, a minimum feature part was used with a 1D deposition technology, such as paste or filament extrusion. A small feature requires a toolpath that includes rapid back and forth motion of the toolhead causing oscillatory vibrations. Measuring acceleration with the fastest travel speeds and smallest predicted feature size showed only subtle vibrations with a maximum acceleration of 0.3 m/s2.4.2 Technology-specific evaluationThe technology-specific customer needs were also evaluated. In order to achieve optimalprocess parameters, an iterative process was used in order to tune the parameters for each individual technology.

In order to achieve optimalprocess parameters, an iterative process was used in order to tune the parameters for each individual technology. First, the process parameters of a certain technology were selected based on similar parameters used in other additive machines. Next, a validation sample with varying features was printed using those settings. After printing a test artifact, the target specifications were reviewed to see if the part met the requirements. If not, the process parameters were adjusted and the process would repeat. Although performance of AM systems can be measured across various metrics, this paper discusses the minimum layer thickness and minimum feature201612size for each technology implemented. Optical microscopy was used to analyze fine features on the test artifacts with high accuracy.Binder JettingBinder jetting layer thickness and feature size are dependent on binder saturation of thepowder.

Enough binder must be jetted to ensure interlayer penetration and overlap between neighboring droplets. While the theoretical minimum feature size is the diameter of a single droplet, the actual minimum feature size is limited by post-processing. The feature must have enough green strength to survive depowdering. Binder jetting parts achieved layer thicknesses as small as 50 m and features as small as 0.47 mm as seen in Table 3 and Figure 11.Table 3. Target specifications and results of minimum layer thickness and minimum feature size for binder jetting.Target SpecificationResultMin. Layer Thickness50 mMin. Feature Size0.47 mmFigure 11. Binder jetting minimum feature size of 0.465 mm.Vat PhotopolymerizationVat photopolymerization utilizing mask projection has the ability to fabricate fine features due to the precise control of digital micromirror devices.

Layer thicknesses as small as 150 m and features as small as 0.12 mm were achieved as shown in Table 4 and Figure 12. With a refined working curve, it will be possible to achieve even smaller layer thicknesses and features by avoiding overexposure.201713Table 4. Target specifications and results of minimum layer thickness and minimum feature size for vat photopolymerization.Target SpecificationResultMin. Layer Thickness150 mMin. Feature Size0.12 mmFigure 12. Vat photopolymerization test artifact with 0.12mm minimum feature.Paste ExtrusionPaste extrusion performance depends on material, pressure, and nozzle size. The multi- tool AM system created a minimum layer thickness of 300 m and minimum feature size of 3 mm, as seen in Table 5 and Figure 13. The figure shows the difference when printing with an applied pressure of 4 psi and 10 psi. Using a less viscous material and a smaller nozzle will allow us to achieve finer features in the future.Table 5.

Using a less viscous material and a smaller nozzle will allow us to achieve finer features in the future.Table 5. Target specifications and results of minimum layer thickness and minimum feature size for paste extrusion.Target SpecificationResultMin. Layer Thickness400 mMin. Feature Size3 mmFigure 13. Paste extrusion minimum feature size test artifacts. The left shows paste extrusion with 4 psi and the right shows paste extrusion with 10 psi of applied pressure.201814Filament ExtrusionFilament extrusion has several key parameters that influence the specifications of interest. Theoretically, the minimum feature should be the diameter of the extrusion nozzle. Although the hot end had a 0.6 mm nozzle on it, this system printed a minimum layer thickness of 500 m and a minimum feature of 1.8 mm as seen in Table 6 and Figure 14. This is due to a limitation of the X and Y linear stages.

This is due to a limitation of the X and Y linear stages. The stages have individual motor controllers that require stopping after a given command. Filament extrusion typically relies on a continuous motion printing process so starting and stopping degrades the quality of the filament track laid. The authors have recognized this limitation and are considering other stages to enhance performance of the system.Table 6. Target specifications and results of minimum layer thickness and minimum feature size for filament extrusion.Target SpecificationResultMin. Layer Thickness500 mMin. Feature Size1.8 mmFigure 14. Filament extrusion minimum feature size of 1.8 mm. This feature size was limited due to XY stages stopping during arc movement.With layer thicknesses and minimum feature sizes determined, the authors created a set of design guidelines to ensure successful manufacturing of parts.

These design guidelines were used to manufacture complex parts with binder jetting and vat photopolymerization as seen in Figure 15.201915Figure 15. Complex parts created via binder jetting and vat photopolymerization.4.3 Multi-material evaluation.Successful implementation of individual technologies lead to the opportunity to combineprocesses. To demonstrate this opportunity, vat photopolymerization was used with paste extrusion. In Figure 16(a) a soft silicone outer shell was extruded with paste followed by an interior rigid grid cured with UV light. In Figure 16(b) a conductive paste was extruded into the interior of a photopolymer cured part to work towards integrated circuits. These demonstrations display the designed functionality that allows the multi-tool AM system to expand AM applications.Figure 16. Multi-material parts fabricated using both vat photopolymerization and paste extrusion.

Multi-material parts fabricated using both vat photopolymerization and paste extrusion. (a) A soft silicone shell was paste extruded with an interior rigid grid created with vat photopolymerization. (b) Conductive paste extruded inside vat photopolymerization parts.The printer has been tested to print with various classes of materials including photopolymers, metals, thermoplastics, polymer powders, and viscous pastes. The individual processes have been tested in order to optimize parameters. More testing with different materials is required to further qualify the system. Additionally, the integration among technologies needs202016to be improved, and the changes between technologies within a part needs to become fully automated. The multi-tool AM system is still being developed and improved, and the authors are confident that with continued advancements, it will soon give researchers the opportunity to revolutionize AM with multi-functional, multi-material parts.5.

ConclusionAll AM technologies share the layer-based approach to fabrication, which allows thepossibility of combining them into one machine. A multi-tool AM system has been invented as the first system to combine five AM technologies in one system. The system has the following unique design features:Toolhead with all technologies in a single assembly  XY movement of toolhead via linear stages   Platform with a reconfigurable build volume  Specific build plates for different technologies  Z movement of build plate via linear actuator  Powder feed chamber for binder jetting  Recoater and overhead heater for binder jetting on linear rails  Sealed build volume for vat photopolymerizationIndividual Z actuation for each toolVat photopolymerization, binder jetting, filament extrusion, and paste extrusion have all been successfully implemented and tested on the system. Material jetting has been designed for and will be implemented into the machine soon.

Material jetting has been designed for and will be implemented into the machine soon. Process parameters for each technology have been tuned in order to meet target specifications. Multi-material parts involving multiple technologies have also been fabricated to demonstrate the functionality of the system. Future work will include improving process parameters to achieve ideal minimum feature sizes, testing new materials, and exploring new multi-material combinations. The multi-tool AM system has successfully expanded the capabilities of AM technologies and has provided the opportunity to explore new AM applications that would not be achievable with single-technology AM systems.6. AcknowledgementsThis project was completed as a Senior Capstone Design Project over two semesters inthe Department of Mechanical Engineering at Virginia Polytechnic Institute and State University.

The authors acknowledge support from and would like to give thanks to the DREAMS lab, the APPLIED lab, the Mechanical Engineering Department, and the Students’202117Engineering Council. The authors would also like to thank Horatio Cowan for his assistance in the APPLIED Lab and LB Fox for his assistance with welding the aluminum build volumes.7.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDESIGN AND FABRICATION OF A BISTABLE UNIT ACTUATOR WITH MULTI-MATERIAL ADDITIVE MANUFACTURINGT. Chen , J. Mueller , K. SheaEngineering Design and Computing Laboratory Department of Mechanical and Process Engineering ETH Zurich, Zurich, 8092, Switzerland ttc · jm · kshea @ ethz . chAbstractBistable systems have two stable equilibrium states around which they can be perturbed. Bistability in a mechanical system requires large deformation, which is commonly accommodated In this paper, the authors leverage the ability of advanced by means of springs or soft joints.

Bistability in a mechanical system requires large deformation, which is commonly accommodated In this paper, the authors leverage the ability of advanced by means of springs or soft joints. Additive Manufacturing (AM) technology that enables printing of materials of varying stiffness values to design and fabricate monolithic bistable unit actuators that maximize the stroke length. A Von Mises Truss (VMT) based bistable structure utilizing snap-through buckling behavior is designed and fabricated. By varying the material stiffness and length of the joints, the authors are able to adjust the activation forces while keeping the geometry constant. 56 specimens are fabricated and tested in tension and compression. Numerical simulation models are constructed and the results are found to correlate well with experimental data.

56 specimens are fabricated and tested in tension and compression. Numerical simulation models are constructed and the results are found to correlate well with experimental data. The required activation force can be increased with a decrease in joint length or an increase in joint material stiffness.IntroductionActive structures have received increasing attention in response to new engineering chal- lenges such as unfamiliar and changing environments [1], but also as a new venue for architectural and artistic expression [2]. The need for such structures arises from two broad categories, one where controlled periodic state change is desirable. Examples include robotic manipulators, active building facades, HVAC systems, and large scale space structures [3]. Other category features structures where the change typically occurs once and must be “reset” after.

Other category features structures where the change typically occurs once and must be “reset” after. Examples in this cat- egory include energy absorbers [4] and gripping and trapping concepts such as an artiﬁcial Venus ﬂytrap [5].As summarized by Korkmaz [1], a typical active structure requires constant power input to remain stable, and must be geometrically accurate and possess structural integrity to carry the designed loads. Multi-material Additive Manufacturing (AM) technologies have been utilized to fabricate active structures. Ge et al. [6] utilized the shape memory effect of printed polymers to “lock-in” pre-stress, and achieve shape change through temperature variation. Tibbits [7] achieved large displacement and conﬁguration changes through swelling of a hydrogel based material also1 2060printed with the same AM technology.

Tibbits [7] achieved large displacement and conﬁguration changes through swelling of a hydrogel based material also1 2060printed with the same AM technology. However, neither approaches offer predictable structural integrity, and lack the premanence and controlability required by a typical active application.Bistability is a buckling-induced mechanical phenomenon that has been increasingly ex- ploited to achieve structural form and functional change [8]. As a bistable structure possesses two equilibrium states, input energy is only needed to transform the state of the structure and not to maintain such a state [9]. Since both the stroke length and the state transitioning force are deﬁned, accurate activation and load bearing capability is inherent in such a structure.In this work, the authors design, fabricate and test a reversible bistable unit actuator that can be fabricated with a multi-material 3D printer monolithically.

The goals of the design are to achieve multiple stable states while maximizing the stroke length and allowing the activation force to be adjustable and predictable. It is uniquely designed for a multi-material fabrication process. As the 3D printer is able to deposit materials with varying Young’s Moduli, the authors tailor the stiffness and the dimension of the joints to allow the bistable actuator to activate with a range of forces without changing its overall geometry. Lastly, the design is fabricated as small as possible to encourage multi-actuator designs that form larger structures.This paper starts with background on bistable mechanisms, the particular AM technology used and joint designs. The following section details an investigation on the printed materials and proposes stress strain relationships for each. Then the geometry of the proposed bistable actuator is described, following with its simulation and validation with physical experiments.

Then the geometry of the proposed bistable actuator is described, following with its simulation and validation with physical experiments. A discussion and conclusion follows.BackgroundThis work combines a bistable mechanism with a multi-material 3D printer to design and fabricate a bistable actuator that achieves the goals set out in the introduction. This section ﬁrst discusses previous fabrication of bistable mechanisms, then provides the theoretical knowledge necessary for the design of a 3D printed multi-material bistable unit actuator. Lastly the impli- cation of the joint fabrication methods are discussed in relation with the multi-material printing technology.Design and FabricationA bistable structure has two stable equilibrium conﬁgurations. After a small force pertur- bation, the structure returns to the initial conﬁguration, similar to a static structure.

After a small force pertur- bation, the structure returns to the initial conﬁguration, similar to a static structure. When the perturbation reaches the magnitude of the activation force, the system suddenly snaps to the sec- ond stable equilibrium. This follows from the sudden large deformation behavior observed when a structure buckles. Bistable phenomenon is exhibited in certain four bar linkages [10], and with limit point buckling based designs such as the Euler buckling beam [11] and the Von Mises Truss (VMT) [12].Euler buckling type bistability consisting of a single beam curved in its ﬁrst buckling mode is popular in literature due to its ease-of-fabrication and large stroke length [13]. Researchers have used such a design as an acoustic damper [14] and energy absorber [15]. The quad-stable structure2 2061proposed by Han et al. [16] focused on the rigidity of the frame constraining the buckling elements.

The quad-stable structure2 2061proposed by Han et al. [16] focused on the rigidity of the frame constraining the buckling elements. Oh and Kota [17] proposed a ﬂat design and stated that single layer monolithic fabrication reduces time and material consumption. Restrepo [15] presented the ﬁrst study of combining the Polyjet printing technology with the fabrication of multi-stable structures, albeit still with a single material. These concepts are adopted and further developed in this work.Bistability with Von Mises TrussesAnother class of bistable actuators exploits the limit point buckling behavior where two un- buckled bars snap to another equilibrium state under a force that is inﬁnitesimally greater than the activation force. Von Mises proposed such a bistable structure in 1923 [12]. A VMT (see Fig. 1a) consists of two truss bars pin jointed at the apex and at the supports.

Von Mises proposed such a bistable structure in 1923 [12]. A VMT (see Fig. 1a) consists of two truss bars pin jointed at the apex and at the supports. VMT exhibits limit-point buckling, where load increase beyond the critical point would change the structural conﬁguration (Fig. 1b). If the vertical load F is applied at the apex and w is the downward displacement from a line drawn between the supports, the relationship between F and w follows Eqn. (1), where E and A are the material stiffness and the cross sectional area of the bars respectively.F = −EA(1 w2 + L2−w01 2 + L2)wWith Eqn. (1), one can plot the load displacement curve (Fig. 1b); the same critical points that exemplify bistability are shown in both ﬁgures (Fig. 1a and 1b). Points U and D are the equilibrium states when the load F is zero. M1 is the critical point at which an inﬁnitesimal increase in load triggers instability.

1a and 1b). Points U and D are the equilibrium states when the load F is zero. M1 is the critical point at which an inﬁnitesimal increase in load triggers instability. With the increase in load, the structure snaps, upon which the VMT essentially mirrors itself along the horizontal axis. With a further increase in load, the bars experience tension and eventually rupture. If the load is displacement controlled, past point M1, the reaction force decreases and eventually reverses orientation past point N and reaches a maximum between point N and D.Joint ConstructionUnlike the Euler bistability design where the whole beam undergoes deformation, the ideal- ized VMTs assume that the truss bars are connected to the bracket with idealized pin joints. In the fabricated design however, the joints must be able to accomodate the large and repeated rotation. Schioler and Pellegrino designed a VMT based device [3] constructed with Nylon injection mold- ing.

Schioler and Pellegrino designed a VMT based device [3] constructed with Nylon injection mold- ing. The members are thinned at the joints and behave as a live hinge. A micro adoption of VMT is used for energy absorption [18], where the material is uniformly ﬂexible.In this work, the authors aim to design and fabricate an actuator such that the stroke to span |w0|/L ratio is as large as possible, and therefore the members in such a design are already at the lower limit of the fabrication resolution. An alternative joint solution is proposed whereby the design utilizes joints with the same cross section but lower stiffness values.

An alternative joint solution is proposed whereby the design utilizes joints with the same cross section but lower stiffness values. This simulates the behavior of a typical joint where the thinning region is replaced by a region where the stiffness is proportionally lower.3 2062(1)1.51M1M20.5x a m FUF0UND0wM10.5N1D1.510.50 wwmax0.51(a) A VMT consisting of two truss bars with geometrical variables labeled(b) Load displacement curve of a typical VMT according to Eqn. (1)Figure 1: Conﬁguration and load displacement curve of a VMT, critical points are labeled.Material StudyThe proposed design relies on a fabrication process that can produce materials of different stiffness values. The Stratasys Objet500 Connex3 achieves this by depositing and mixing liquid photopolymers from an array of nozzles in a layer by layer fabrication process.

The Stratasys Objet500 Connex3 achieves this by depositing and mixing liquid photopolymers from an array of nozzles in a layer by layer fabrication process. UV light cures the recently printed layer before the print platform lowers and the next layer is printed. This allows the different materials to be assigned to any region of a design.With this 3D printer, available stiffness values are not a continuous variable. Rather the printer allows the mixing of two or more base materials in 12 discrete steps, yielding 14 different material combinations for each pair. Of particular interest are the mixtures between TangoBlack+ (TB+), an elastomer-like material, and VeroWhite+ (VW+), a rigid plastic to create 12 intermediate digital materials (DMs) of varying stiffness values.The material manufacturer provides the Young’s Modulus and fracture strain for the seven DMs on the rigid end of the spectrum but not for the ﬂexible ones. Stankovi´c et al.

Stankovi´c et al. [19] pre- sented the Young’s Modulus of all 14 materials, showing large deviations from the manufacturer’s speciﬁcation and the stress-strain relationships were not included. The deviations were attributed partly to unknown test conditions [20]. As bistability entails large deﬂection [21] and ﬁnite strain behavior of the elastomers, quasi-static tensile tests are conducted under constant environmental conditions in this study. The results are used in the simulation models that predict the behavior of the unit actuators.MethodThis section describes the proposed bistable actuator and shows that the actuator achieves the goals described in the introduction. Namely, the proposed bistable actuator possesses two4 2063equilibrium states and is reversible; that the actuator can be fabricated monolithically with a 3D printer and have predictable activation force that can be tailored through joint length and material change.

Lastly, the stroke length is maximized while keeping the design as small as possible.To demonstrate bistability as well as predictability of the activation force, ﬁrst a material char- acterization is performed to determine the mechanical properties of the printer materials. Second, a parametric study is conducted both with numerical simulation and with physical experiments on 56 specimens of varying joint length and joint material.Unit Actuator DesignThe proposed bistable actuator shown in Fig. 2 originates from the VMT scheme. The phys- ical realization requires rigidity laterally and out of plane. As suggested in [3] and [22], the side- way movement is restricted by stacking two VMTs vertically, and the out-of-plane movement is restricted by increasing the depth of the entire actuator. In using such an actuator as a unit, and to maximize the stroke length, the design must be self-contained and small.

In using such an actuator as a unit, and to maximize the stroke length, the design must be self-contained and small. This places an empha- sis on the design of the ﬂexible joints critical to VMTs. It was stated that compared to an Euler buckling based bistable mechanism, the stroke distance provided by VMTs is much smaller [3]. To overcome this limitation, the proposed actuator provides a |w0|/L ratio of 1 where w0 and L are the width and the height of the actuator respectively.

The bracket, the pin and the center portion of the truss bars are fabricated with VW+, and are much stiffer than the joints indicated5 2064by the line-hatched regions. The design is dimensioned so that the smallest elements coincide with the precision of the fabrication method. Note that the bracket cannot be assumed to be rigid as its ﬂexibility partly allows the actuator to function. All variables are listed in Table 1. The joint geometry is created such that it is perpendicular to both the bracket and the pin to ensure the contact area remains constant and does not depend on the inclination angle, α. The cross-hatched regions indicate the additional parts included to facilitate tensile testing. When removed, one notes that the bottom of the bracket has the same geometry as the pin.

The cross-hatched regions indicate the additional parts included to facilitate tensile testing. When removed, one notes that the bottom of the bracket has the same geometry as the pin. Therefore, when many actuators are stacked together in series, the total extended length nearly doubles the retracted length.The design is assumed to be rigid in the out-of-plane direction and has a thickness of 4 mm. The angle of inclination of the truss bars, α = tan(w0/L), is set to 45°. The stroke length, 2w0, is dependent on the inclination angle. The minimum spacing of the fabrication process k dictates how close two features may be placed without being fused together. The extrusion of length h is added to the design for simulation and testing purposes to maintain a predictable load condition across all specimens. It is important to note that the extrusions on the pin and on the brackets are in line with each other and therefore no applied moment is generated.

It is important to note that the extrusions on the pin and on the brackets are in line with each other and therefore no applied moment is generated. The line-shaded region indicates the location of the joint and a change in material stiffness. Parameter l deﬁnes the length of the joint and is speciﬁed in the parameter study section. Note that changing l does not change the overall geometry, i.e. if the joint lengthens, the truss itself shortens.Printer Material CharacterizationTo adequately simulate, as a part of the parametric study, the large displacement and strain necessitated by the VMT mechanics, a complete stress-strain model for each joint material must be provided. With the Connex printer, between TB+ and VW+, 12 DMs can be deposited within one print. The last two digits of each material name is its Shore hardness values (Shore A for the six DMs in the ﬁrst group, and Shore D for the second). The complete set of materials is shown in Fig.

The last two digits of each material name is its Shore hardness values (Shore A for the six DMs in the ﬁrst group, and Shore D for the second). The complete set of materials is shown in Fig. 3.Base MaterialDigital MaterialG roup 1 G roup 2TB+ VW+F9840 R8505F9850 R8510F9860 R8515F9870 R8520F9885 R8525F9895 R8530Increase in stiffnessFigure 3: Material table with labels, grouping and base/digital distinction shown. Direction of stiffness increase is indicated.The following describes the tensile testing procedure that is performed for each material in Fig. 3. Type V geometry of the ASTM D638-10 test standard for tensile testing of plastics is adopted for all specimens. Three identical test specimens are printed for each material under an ambient temperature of 21 ◦C. To eliminate the impact of uncontrolled factors on the mechanical properties [20], each is printed on a separate tray. Aside from the material combination, all other factors are held constant.

To eliminate the impact of uncontrolled factors on the mechanical properties [20], each is printed on a separate tray. Aside from the material combination, all other factors are held constant. Tensile tests are conducted on an Instron ElectroPuls 3000 with either a 100 N or a 5000 N load cell, depending on the material tested. Under displacement control, a testing speed of 20 mm min−1 is set and the specimens are elongated until fracture.6 2065Parametric Design StudyWith both the actuator geometry and the material properties deﬁned, a study can be conducted to assess the effect of the joint on the overall behavior and the activation forces. As discussed in the actuator design section, the varying parameters are the geometry and the material of the VMT joints. The printed position, retracted or extended, of the actuator during fabrication is included as an additional parameter.

The printed position, retracted or extended, of the actuator during fabrication is included as an additional parameter. Along with four different joint lengths and seven different joint materials, the parametric variations lead to a total of 56 models. These models are compared with the analyt- ical solution, simulated numerically, and are fabricated. Quasi-static load testing is conducted on the fabricated specimens. Figure 4 shows the different parameters and the corresponding geome- try and printed specimens. It is hypothesized that decreases in the joint length or increases in the material stiffness of the joint will increase the activation force.A. Joint LengthB. State at Fabrication1.000.750.500.25C.

Joint LengthB. State at Fabrication1.000.750.500.25C. Joint MaterialTB+F9840 F9850 F9860F9870 F9885 F9895RetractedExtendedFigure 4: The design parameters including joint length, actuator state at fabrication and joint ma- terial investigated in the parametric study are shown by focusing on one jointAnalytical Solution The analytical solution in terms of Eqn. (1) is used for comparison. Two shortcomings are noted in the analytic model. One, a VMT assumes the joints are idealized pins, where as in the design presented in this work, joints have ﬁnite geometry and stiffness. Considering this, the authors propose to substitute E with a modiﬁed Young’s Modulus assuming each truss is a series of connected springs of varying stiffnesses, i.e.

Considering this, the authors propose to substitute E with a modiﬁed Young’s Modulus assuming each truss is a series of connected springs of varying stiffnesses, i.e. Ej represents the stiffness of the joint material.7 2066EMod =2 1 Ej1 + 1EVW+The second shortcoming is that the bracket is assumed to be rigid in the analytical solution, whereas in reality, they bend as the pin is extended. One potential way to incorporate this effect in the theoretical equation is to add a linaer spring in between the supports and assume that the supports are a pin and a roller, and not fully ﬁxed. This will be investigated in a future study.Numerical Simulation A ﬁnite element based simulation is performed to predict the behavior of a unit actuator. With the intent to reduce computational effort rather than a formal ﬁnite element comparison as done by Cherry [22], two sets of simulations are implemented.

With the intent to reduce computational effort rather than a formal ﬁnite element comparison as done by Cherry [22], two sets of simulations are implemented. The ﬁrst aims to replicate both the form and the boundary conditions of a physical actuator (Fig. 5A). As such, brick elements are used to mesh the volume. A ﬁnite displacement of 10 mm min−1 is applied to surface A of Fig. 2; surface B is pin jointed. Symmetry conditions are imposed on surfaces C and D to reduce computational effort. Material properties summarized in Table 2 are incorporated into each analysis. Since this is a displacement based simulation, a physically and geometrically non-linear analysis is performed without arc-length control.The second set of simulations assumes that the depth and the width of each section can be approximated as that of a rectangular beam. With that, the model is reduced to a frame and meshed with quadratic Timochenko beam elements.

With that, the model is reduced to a frame and meshed with quadratic Timochenko beam elements. The beams are placed at the centerline of the solid geometry. Where there is a connecting point between two members, a rigid link is placed to eliminate artiﬁcial ﬂexibility (Fig. 5B). The boundary conditions and the material model remain the same as the previous simulation model. With the beam model, each of the parametric variations are numerically simulated.The focus of this study is not on the eventual failure of the joints. Therefore, the displace- ment load is restricted between 0 and the stroke length. No failure criterion is set, nor is there a contact condition between the pin and the bracket in the case of compressive displacement with the extended specimens. The solid model is used as a benchmark of the parametric study that uses the beam model. As such, a random case of l = 0.75 mm and FLX9860 as the joint material is selected as the benchmark model.

The solid model is used as a benchmark of the parametric study that uses the beam model. As such, a random case of l = 0.75 mm and FLX9860 as the joint material is selected as the benchmark model. Both retracted and extended states are simulated, resulting in two simulation models.Experimental Validation The aforementioned models are fabricated on the multi-material 3D printer. The surface ﬁnishing is set to glossy to ensure that the specimens are not encased by support material. The support material does not dissolve with a solution and must be washed away with a water jet. With delicate specimens, the effect of the water jet must be avoided. The Instron appratus is used to conduct both tensile and compressive testing. All specimens are tested on the day they are fabricated and no state change occurs prior to testing.

The Instron appratus is used to conduct both tensile and compressive testing. All specimens are tested on the day they are fabricated and no state change occurs prior to testing. To ensure that the boundary conditions remain invariant across all models, a set of two customized grippers are printed to hold the specimens by the protrusions dimensioned h shown in Fig. 2. The grippers themselves remain clamped to the Instron at a pressure of 0.5 MPa. For the retracted specimens, a tensile displacement of 10 mm min−1 is set and the specimens are elongated until rupture. For the extended specimens,8 2067(2)Figure 5: The Solid (A) and Beam (B) FEM model with material and interfaces shown. (C) The benchmark actuator printed in the retracted state, and extended through an applied tensile force (D). The experimental setup showing the grippers (E).a compressive displacement of equal magnitude is applied.

The experimental setup showing the grippers (E).a compressive displacement of equal magnitude is applied. Both the actual displacement and the reaction force are recorded for each specimen.ResultsThis section presents both the result from material characterization and that from the para- metric study. From the material characterization, ﬁrst the force displacement curves of each of the 14 materials are shown. Then a multi-linear elastoplastic constitutive model is constructed for each material to be used in the numerical simulation. From the parametric study, the validity of the theoretical model and of the two simulation models are compared using the benchmark specimen (i.e. joint material - F9860 and joint length - 0.75 mm). Then results from the beam model are compared to the physical experiments. Lastly, the activation force required by each specimen is investigated.Material CharacterizationFrom the exhibited behavior (Fig.

Lastly, the activation force required by each specimen is investigated.Material CharacterizationFrom the exhibited behavior (Fig. 6), the two dominating materials of the mixing process, TB+ and VW+ can be distinguished. Qualitatively, materials printed with TB+ as the base fail slowly, whereas specimens printed with VW+ as the base exhibit necking and fail subsequently from brittle fracture. From material TB+ to VW+, the strength gradually increases, whereas the fracture strain decreases. In the ﬁrst group (TB+ to F9895), the fracture strain ranges from 30 % to 174 % and ultimate strength ranges from 0.86 MPa to 10.6 MPa. In the second group (R8530 to VW+), the fracture strain ranges from 8 % to 14 % and ultimate strength ranges from 54 MPa to 69 MPa.

In the second group (R8530 to VW+), the fracture strain ranges from 8 % to 14 % and ultimate strength ranges from 54 MPa to 69 MPa. In general, the observed properties of the DMs can be grouped around the properties of9 2068the two base materials, creating a gap in printable material properties in between.(a) TangoBlack+ dominating group(b) VeroWhite+ dominating groupFigure 6: Experimental results showing stress strain curve of each printer material.This bi-modal distribution of the material properties validates the assumption that the mixed materials’ properties depend heavily on the base material used. Attention is focused on the TB+ dominating group where the allowable strains are deemed sufﬁcient to accommodate the large deformation induced by the snapping behavior and the range of allowable stresses are sufﬁcient to deliver the desired activation forces.

To simplify the material behavior, each of the 7 TB+ based materials is assumed to be able to be adequately modeled by a tri-linear stress strain relationship. The elastic Young’s Modulus of each material can be calculated by dividing the ﬁrst stress by the ﬁrst strain of that material.

F9840 and F9850 can be described bi-linearly; the strain value at the ﬁrst control point of these materials, 200 %, are much higher than that of the stiffer materials. F9860 to F9895 are described tri-linearly, the second strain10 2069point is set at 5 % to describe the softening effect and the third strain point is set at rupture. The behavior of VW+ is distinct from the rest as it belongs to the rigid material group and follows the behavior shown in Fig. 6b.Analytical and Numerical ModelsFor the benchmark case, with the geometric data from Table 1 and the material data from Table 2, we obtain the idealized curve in Fig. 7. In the same ﬁgure, the force displacement curve of both initial states and of both numerical models are shown. From the ﬁgure, bistable behavior can be seen in all models. However, the analytic model presents a symmetric bistable behavior, whereas the numerical models of the actual design show an asymmetric behavior.

However, the analytic model presents a symmetric bistable behavior, whereas the numerical models of the actual design show an asymmetric behavior. The stable equilibrium , and the unstable ones with ×. Qualitatively similar results have positions are indicated with been presented in literature [22]. The actual stroke length is shorter than intended as the force is not zero at the fully extended state (w = ±4.5 mm). Rather, the curve crosses the zero force line at approximately 2.5 mm for the retracted initial position and −3.0 mm for the extended one, giving the actuator a stroke length of 7 mm and 7.5 mm respectively. Comparing between the solid and the beam numerical models of the retracted state, the behavior is similar. This is in contrast with [22] where it was concluded that the 3D model offered closer agreement due to out-of-plane effects not captured with a 2D model.

This is in contrast with [22] where it was concluded that the 3D model offered closer agreement due to out-of-plane effects not captured with a 2D model. With respect to the analytic model, the differences in the activation force is 8.8 % for the beam model and 20 % for the solid model. The same observation is made in the extended state, except the differences are greater for the activation force, 55 % and 89 % for the beam and the solid model respectively.(cid:2)Figure 7: Plots comparing the analytical solution, and simulation featuring solid or beam elements for both retracted and extended initial positions11 2070Beam Model and Experiment ResultsAs it is difﬁcult to ﬁguratively show the results of all 56 models, plots are made in which only one parameter varies. First, the effect of joint length is investigated by keeping the joint material ﬁxed (Fig. 8). The lengths studied are 1.0, 0.75, 0.50, 0.25 mm (Fig.

First, the effect of joint length is investigated by keeping the joint material ﬁxed (Fig. 8). The lengths studied are 1.0, 0.75, 0.50, 0.25 mm (Fig. 4), however during the experiment l = 0.25 mm consistently ruptures while undergoing state change. Therefore, designs with such a length are excluded.(a) Initial position: Retracted(b) Initial position: ExtendedFigure 8: Force-displacement behavior from experiments and simulation with ﬁxed joint material stiffness and varying joint length, * denotes the benchmark conﬁgurationFor the retracted initial state, the behavior between experimental data and simulation show remarkable similarity, with a root-mean-square deviation(RMSD) equal to 0.156, 0.115, 0.121 for joint lengths 1.0, 0.75, 0.50 mm respectively. With the initial state being extended (Fig. 8b) the general observations are the same as above. The RMSD of joint lengths 1.0, 0.75, 0.50 mm are 0.402, 0.425, 0.249 respectively.The second set of ﬁgures (Fig.

8b) the general observations are the same as above. The RMSD of joint lengths 1.0, 0.75, 0.50 mm are 0.402, 0.425, 0.249 respectively.The second set of ﬁgures (Fig. 9) keep the length at l = 0.75 mm while varying the joint mate- rial from TB+ to F9895. Predictable bistable behavior is observed as the activation force increases with increasing joint stiffness. The discrepancy between simulation and experiment increases at the extremes of joint material stiffness and lessens at the center, with RMSD equal to 0.213, 0.174, 0.143, 0.115, 0.227, 0.479, 0.703 for the retracted initial position and 0.682, 0.649, 0.533, 0.235, 0.363, 0.562, 0.851 for the extended case.Figures 8 and 9 present results from 18 of the 56 specimens (note that designs where the joint length is 0.25 mm are omitted so the number of useable specimens is 42). The same qualitative observations can be made with the remaining specimens and are omitted here.

The same qualitative observations can be made with the remaining specimens and are omitted here. The important quantitative results are presented below.Activation ForceAs the unit actuator intends to provide varying activation force by changing material param- eters, the spectrum of achievable activation forces is plotted in Fig. 10 for both simulation and12 2071(a) Initial position: Retracted(b) Initial position: ExtendedFigure 9: Force-displacement behavior from experiments and simulation with ﬁxed joint length and varying joint material stiffness, * denotes the benchmark conﬁgurationexperimental results. In the horizontal axis are the specimens numbered according to Table 3. On the vertical axis is the peak force attained before snap-through occurs in each specimen. Both the retracted and the extended states are shown.Table 3: Specimen num- ber and corresponding joint length and materialm m 0 . 1m m 5 7 . 0m m 0 5.

Both the retracted and the extended states are shown.Table 3: Specimen num- ber and corresponding joint length and materialm m 0 . 1m m 5 7 . 0m m 0 5 . 03 2 1 TB+ 6 7 5 F9840 F9850 10 11 9 F9860 13 14 15 F9870 17 18 19 F9885 21 22 23 F9895 25 26 27Figure 10: Activation force of each specimen - simulation and experimental resultsWith the retracted specimens, an activation force ranging from 0.50 N to 5.0 N can be achieved. With the extended specimens, the activation force range is 0.25 to 3.5 N. Simulation of the acti- vation force is able to predict experimental results with conﬁdence, with the highest accuracy in the mid-range; simulation underestimates the activation force of actuator units using softer joint materials and overestimates the activation force of actuator units using stiffer joint materials.13 2072DiscussionA unit actuator design leveraging different material stiffness fabricated using a multi-material 3D printer is proposed.

The design is built on the idealized Von Mises Truss model where two truss members are connected by three pins, and is an improvement over the design proposed by Schioler and Pellegrino [3] in terms of overall size and stroke length to half span ratio. The stroke length over span of 100 % is signiﬁcantly larger than comparable designs, Restrepo et al. [15] achieved a stroke ratio of 16 % with an Euler buckling based bistability. Schioler and Pellegrino’s VMT based design has a stroke ratio of 29 % [3], and [23] using a similar design achieved 24 %. It is worth noting that the overall geometry of an Euler buckling based design encourages stacking as shown by Restrepo et al. [15].The study sets out to design an actuator that is able to offer a range of activation forces without changing its overall geometry, and the proposed design demonstrates this property in the results section.

Either by shortening the length of the joints, which lengthens the actual truss, or by increasing the stiffness of the joint material, the activation force increases predictably (Fig. 10). These parametric changes are made possible through a 3D printer that is able to deposit material of varying stiffness in a single fabrication. The design is dimensioned so that it can be stacked vertically, where the bracket of the actuator above serves at the pin of the actuator below.A number of shortcomings are noted as well. The discrepancies between the theory and the simulation models lie largely in three aspects, completeness of the material characterization, the asymmetrical buckling behavior and accuracy of the simulation results (i.e. differences in the forces between the curves in Fig. 7). Each is discussed in turn. One, each digital material undergoes tensile testing only; the compressive behavior is assumed to be identical. Further, only one build orientation is considered.

7). Each is discussed in turn. One, each digital material undergoes tensile testing only; the compressive behavior is assumed to be identical. Further, only one build orientation is considered. The printed plastics possess viscoelastic properties that are not addressed in material testing or in simulation. Two, the asymmetrical behavior may be caused by the following. The bracket of the actuator cannot be assumed rigid as in the idealized VMT as one may note signiﬁcant bending in Fig. 11. This leads to signiﬁcantly different structural behavior between the top and the bottom trusses. Three, the lack of accuracy may be explained as follows. As shown in Fig. 5C, the printed specimens do not perfectly replicate the design intent. The measured dimensions of each member is within ±0.1 mm of the design. Due to the glossy ﬁnishing, the top surface is not ﬂat and the edges appear to be ﬁlleted. At the material interface, signiﬁcant blending of the two materials can be observed.

Due to the glossy ﬁnishing, the top surface is not ﬂat and the edges appear to be ﬁlleted. At the material interface, signiﬁcant blending of the two materials can be observed. These defects do not usually alter the performance of a large product, however they may have a pronounced effect at this scale.Other important limitations exist primarily in the material behavior. One, all the material testing specimens are fabricated along the x-axis, whereas the unit actuators consist of members printed in all angles between x and y. This has an inﬂuence on the simulation results as the print materials are not isotropic [24]. Two, it is observed that the TB+ based materials degrade over time similar to natural rubber. This effect is controlled as the experiments are conducted on the day of fabrication. Fatigue is a major concern with such structures [21]. Its effect is controlled in this study as no specimens underwent state change prior to testing.

Fatigue is a major concern with such structures [21]. Its effect is controlled in this study as no specimens underwent state change prior to testing. Both time dependent effects and fatigue should be included in future studies to assess the robustness of the actuators.Since this study focuses primarily on designing actuator conﬁgurations to provide the maxi-14 2073Figure 11: Deformed shape and stress distribution of an actuator in different stages of activation, note the signiﬁcant bending in the bracket when the truss bars are close to being horizontal.mum stroke length and a spectrum of activation forces, a further study is planned to vary geometric and material parameters to tailor to desired overall behavior. In particular by varying the stiffness and length of individual joints, one can plan a predictable and repeatable activation sequence.The next logical step in this study is to incorporate such a unit actuator in designs that require many equilibrium states.

Each unit may be thought of as a binary switch, and the number of stable states equal to 2n where n is the number of units in the structure. The ﬁrst step is to compound the activation forces or stroke distance in simple ways. Assuming that each actuator behaves linearly up to the point of activation, by placing two in parallel, the structure would double the force necessary to cause a state change. If one places two actuators in series, the stroke length would double. The design of multi-stable hierarchical structures based on the unit actuator proposed in this paper is presented in Chen and Shea [25].ConclusionA bistable actuator is designed for a multi-material 3D printer and its behavior is investigated in this work. Adapting from the Von Mises Truss model where two truss members are connected by three pin joints, the idealized joints are fabricated with materials that are much less stiff than the part material.

Adapting from the Von Mises Truss model where two truss members are connected by three pin joints, the idealized joints are fabricated with materials that are much less stiff than the part material. The actuators are fabricated with a 3D printer that is able to deposit material of different stiffness values, and designed to be stackable, thereby maximizing the stroke length to half span ratio. A reduced simulation model using beam ﬁnite elements is able to predict experi- mental results with close agreement. Tensile testing is conducted and an experimental stress-strain curve is obtained for each material. With such constitutive models, a parametric design study of the joint length and material is done both through simulation and experimentation. By varying the joint length from 0.25 to 1.0 mm and joint material stiffness from 0.485 to 375 MPa, the actuator requires different activation forces while maintaining constant geometry and stroke length.

sign And Fabrication Of Components With Optimized Lattice MicrostructuresVito R. Gervasi (Co-PI) and Douglas C. Stahl (PI) the Milwaukee School of Engineering, Milwaukee, WisconsinABSTRACT:The design and fabrication of components with optimized lattice microstructures is a new approach to creating lightweight high-performance objects. This paper introduces a unique and complete integration of design and fabrication leading to the creation of structural components with complex composite microstructures. Rather than a solid cast component with optimized outer shape this new approach leads to a component with an inner skeleton (cid:150) or microstructure (cid:150) maximizing one or more properties such as the stiffness-to-weight ratio. Three dimensional gradient materials are a natural outcome of this approach. An introduction to the design optimization and hybrid fabrication approach will be provided in addition to research progress and challenges through Spring 2004.1.

An introduction to the design optimization and hybrid fabrication approach will be provided in addition to research progress and challenges through Spring 2004.1. INTRODUCTIONMany biological composites have properties (strength, toughness, etc.) that are truly remarkable when compared with the properties of the constituents. Mollusk shells, for example, have high flexural strength and good fracture toughness, but they are made almost entirely of calcium carbonate. As Bond et al. [2] pointed out, imagine the possibilities if one could design a comparable composite based on high performance constituent materials. Results of the biological or evolutionary (cid:147)design and fabrication process(cid:148) are no less impressive when one considers whole components or systems. Srinivasan et al.

Results of the biological or evolutionary (cid:147)design and fabrication process(cid:148) are no less impressive when one considers whole components or systems. Srinivasan et al. [8] noted that (cid:147)biological structural systems do not distinguish between [design of] materials and structures.(cid:148) In other words, the large-scale structure is optimized, and the material from which it is made is optimized.One goal of this project is to develop a process to integrate optimization of a structural component(cid:146)s shape and topology with optimization of the composite material in it, by treating the component(cid:146)s inner skeleton as part of the design domain.

Rather than a solid cast component with optimized outer shape, for example, one could produce a component with an inner skeleton (cid:150) or microstructure (cid:150) designed to maximize the stiffness-to-weight ratio.The second goal of this project is to develop hybrid fabrication techniques to manufacture the components with complex microstructures that result from optimized design. Solid Freeform Fabrication (SFF) has emerged as a powerful method for producing complex three-dimensional objects, so the concept of fabricating components with complex microstructures has become realistic. The current research focuses on hybrid fabrication techniques, which combine SFF with other fabrication steps such as casting and machining.

The current research focuses on hybrid fabrication techniques, which combine SFF with other fabrication steps such as casting and machining. The main benefit of hybrid fabrication is that it permits one to use a wide range of materials, adding each under its own ideal conditions.The present paper describes progress made during the first year and a half of the projectand some of the challenges currently facing the project team.8382. PROGRESS AND CHALLENGES IN THE DESIGN OF OPTIMIZED MICROSTRUCTURES The objective of structural optimization is usually to maximize the stiffness of a structure without adding weight. This is expressed mathematically as a minimization of strain energy subject to various constraints. Shape and size optimization, in which the connectivity or relationship between various parts of the object is pre-established, are becoming standard design tools for parts manufactured by casting, forging, and stamping.

Topology optimization is a more general procedure; it can address connectivity and the placement of voids, and also the arrangement of material of varying density. One critique of some types of topology optimization has been that the presence of regions with intermediate densities is unrealistic. The premise of this project is that regions of varying density are not unrealistic at all (cid:150) the manufacturing capabilities of SFF and Hybrid Fabrication make density variation possible. Preliminary work has focused on testing concepts for the optimization process and exploring strategies for transferring the design information to the SFF machines. The structural components considered to date are limited to two-dimensional objects subjected to single in- plane load cases.

The structural components considered to date are limited to two-dimensional objects subjected to single in- plane load cases. All computer programs used for optimization and processing of design data have been written by the PI and an undergraduate research assistant.The optimization procedure implemented for preliminary work follows in general the SIMP or Solid Isotropic Material with Penalization optimization [7]. The optimized design program with its finite element core is written in FORTRAN with efficient sparse matrix manipulation functions [4]. A graphic record of the optimization process is shown in Fig 1. This planar structural component is supported along the lower left edge, and subjected to concentrated loads at the upper right corner directed downward and rightward. The planar area was divided into a 100 by 100 grid of four-node plane elements. Initially all elements are given the same density.

The planar area was divided into a 100 by 100 grid of four-node plane elements. Initially all elements are given the same density. Material density was treated as the variable, and the structure(cid:146)s overall weight was held constant while the stiffness was maximized. The optimization process was permitted to run through only a few iterations to create a design that includes a wide range of densities. TheFigure 1. Topology optimization: the cantilever is fixed at the lower left corner; vertical and horizontal loads are placed at the upper right corner. High density regions are darker.problem was re-run with a coarser mesh, with the resulting density map shown in Fig. 2.839Interpreting the density map to describe a physical object is the next step in the process. The design information was post-processed in two ways. First, the density map was interpreted as a thickness map. The resulting component, the upper right corner of which is shown in Fig.

The design information was post-processed in two ways. First, the density map was interpreted as a thickness map. The resulting component, the upper right corner of which is shown in Fig. 3, could be molded, forged, or made by many other traditional fabrication techniques. Second, the density map was interpreted literally, and a microstructure was manipulated to provide the desired density. The microstructural unit cell is a rods are lattice. The three section, and density rectangular determines the cross-sectional dimensions of the rods to the overall dimensions of the unit cell. The resulting component, shown partially in Fig. 4, has uniform thickness set arbitrarily at four times the unit cell dimension.rod orthogonal in cross theratio ofFigure 2. Density map from optimized design program.One result of the design process not incorporated into the object shown in Fig. 4 is the map of principal stress directions.

Density map from optimized design program.One result of the design process not incorporated into the object shown in Fig. 4 is the map of principal stress directions. While imagining a procedure to re-orient the lattice elements of Fig. 4 to conform to principal stress direction, one can see that simply re-orienting the lattice units inside the existing array of rectangular cells would not work, as adjacent lattice units would no longer be aligned with one another. An entirely new array of cells (cid:150) with the array aligned with principal stresses so the lattice structure will also be aligned (cid:150) is required. This is analogous to a central step in adaptive finite element work, where a simple mesh is used for a preliminary analysis and then a better mesh, automatically generated from the preliminary results, is used for a refined analysis.Figure 3.

Interpretation of density as thickness (upper right corner of Fig 2).A second program was written by the PI to generate a new mesh of quadrilateral cells based on the optimization results. Mesh generation for this application uses, in addition to the density and principal stress direction maps, a map of principal to stress direction gradient which determine element size. The maps are searched numerous times during mesh generation; implementation of an (cid:147)alternating digital tree(cid:148) algorithm [3] to organize the information in these maps is essential for efficient searches. Automatic mesh generation follows, in general, the procedure described by Zhu et al. [10]. The resulting mesh, for the example of Fig. 2, is shown as Fig. 5. The mesh consists only of quadrilaterals, the element boundaries follow the principalFigure 4. Microstructural interpretation of density (upper right corner of Fig 2).is used840Figure 5. Automatic generation of new mesh.Figure 6.

Microstructural interpretation of density (upper right corner of Fig 2).is used840Figure 5. Automatic generation of new mesh.Figure 6. Resulting lattice microstructure.stress directions (with deviations due to the constraints of creating and smoothing the quadrilaterals), and regions with sharp changes in principal stress direction have smaller elements.A third program for manipulating the output of the optimized design program was written in C++ by an Undergraduate Research Assistant. This program requires as input the mesh shown in Fig. 5, along with information about the material density at each face of each cell. The program produces the three-rod lattice in each cell, with each rod having a regular polygon cross section and tapered to interpolate density. The output is an STL file which is the standard input to the SFF machines. The STL file, graphically represented in Fig. 6, describes the surface area of the three dimensional object as a set of triangles.

The STL file, graphically represented in Fig. 6, describes the surface area of the three dimensional object as a set of triangles. Additional discussion of the use of the STL file was presented in [9].The difference between the two interpretations of data from the optimization procedure shown in Figs. 3 and 6 is critical. The first interpretation is consistent with the current state of the art in industrial design. The varying density demanded by optimized design is interpreted as varying thickness for a nominally two-dimensional object. Because this interpretation uses the third dimension to create the effect of varying density, it could not be used to make objects with density varying in three dimensions. The second interpretation uses the microstructure to control density and orientation (and therefore the effective material properties).

The second interpretation uses the microstructure to control density and orientation (and therefore the effective material properties). The third dimension is still available, so this interpretation could be used to make structural components with microstructural gradients of density, orientation, stiffness, or other effective property in three dimensions.Current challenges in designHaving completed the algorithm development and programming for a first draft of the entire design process during the project(cid:146)s first year, work has turned to improving the process and making it more general. During the 2003-4 academic year attention has turned to these areas with much progress to date: • The homogenization approach to topology optimization [1,6] is being implemented to841recognize the anisotropic nature of the final design during the optimization process.

Instead of a simple density map, this will produce (for two-dimensional problems) maps of lattice rod dimensions for local x and y directions.The MSOE Rapid Prototyping Center announced that the model of Fig. 6 broke the record for the most errors in an STL file. Fortunately the staff was able to get the Selective Laser Sintering machine to build the part. To ensure portability of designs to other SFF machines, we are developing algorithms and programming to create (cid:147)error-free(cid:148) STL files.Design for multiple load cases. • Optimization of three-dimensional objects.3. PROGRESS AND CHALLENGES IN HYBRID FABRICATIONHybrid fabrication is a new approach for making multi-phase composites, including functionally graded materials, by combining the most advantageous aspects of SFF and conventional fabrication techniques.

Here the critical geometric feature of the composite is intertwining two or more three-dimensional lattices, thus each phase of the composite is continuous and separate from the other phases. SFF is used to create the initial framework for the intertwined lattices, and then the solid materials for each phase are added as separate steps as shown in Fig. 7 [5]. Hybrid synthesis of intertwined lattices avoids problems that can occur when multiple materials are directly formed in a SFF process by combining two or more powder types within a melt-pool. One such problem is the creation of multiple metallurgical grain microstructures and porosity. Another challenge is dealing with varied material melting points, delivery systems, and buoyancy of ceramic reinforcements.

Another challenge is dealing with varied material melting points, delivery systems, and buoyancy of ceramic reinforcements. By separating the materials into lattices and creating one lattice phase at a time using hybrid fabrication, the metallurgical challenges are avoided; each material can be processed under its own ideal conditions.Progress in hybrid fabricationthere are no standard hybrid synthesis procedures for generating complex lattice structures with features smaller than 1 mm, so a starting point for this project was to consider a large number of SFF-based hybrid casting (cid:147)paths(cid:148) and combinations of steps. Metal casting techniques such as centrifugal casting were experimented with, leading to concerns and questions about quantifying the quality of castings and limitations of select casting processes.

Metal casting techniques such as centrifugal casting were experimented with, leading to concerns and questions about quantifying the quality of castings and limitations of select casting processes. With no standards for comparing the potential of hybrid fabrication paths a standard benchmarking system was developed.Currently,Evaluation of a number of hybrid paths, on paper and experimentally, has led to a list of hybrid fabrication requirements as follows: the complete transferal of the expendable mold material around (and into) very fine features of the expendable pattern without trapping gas or forming voids; complete removal of the expendable pattern without damaging the mold; ceramic mold firing without cracking,Figure 7.

Example Hybrid Fabrication Process842sagging, or breakage; pattern removal without residue in (cid:147)unvented dead-ends(cid:148) and fine mold features; complete mold fill with metal without damaging or eroding the mold geometry; complete mold filling with metal before freezing; overcoming surface tension of molten metal to fill features finer than 1 mm; elimination of gas defects in casting due to trapped gasses; control of the direction of solidification to reduce solidification shrinkage defects; prevention of buoyancy force damage on delicate, high aspect ratio, features.The development of a new hybrid fabrication process has begun, with preliminary results producing castings with features approaching 0.05 mm. The process is similar to that illustrated schematically in Fig. 7. This approach has good potential to meet the aforementioned requirements including resolution capabilities and ability to be processed in a vacuum.

7. This approach has good potential to meet the aforementioned requirements including resolution capabilities and ability to be processed in a vacuum. Several breakthroughs in ceramic processing, pattern removal and metal casting have resulted in the casting of very fine features, such as those of the fractal tree shown in Fig. 8.Figure 8. Finely detailed casting (cid:150) a result of hybrid fabrication..To evaluate potential hybrid fabrication processes, the effect of each processing variable on the quality of the finished part must be assessed. To this, several benchmark parts were developed. An example is shown at the top of Fig. 9; this part examines minimum feature size of lattice rods. Fig. 9 also shows results of a that hybrid process similar to shown features for representing rods on the final metal casting.accomplishin Fig.

Fig. 9 also shows results of a that hybrid process similar to shown features for representing rods on the final metal casting.accomplishin Fig. 7Current challenges fabricationin hybridWith one practical hybrid fabrication path selected and a method to measure capability and improvement the research focus has been directed to a number of new challenges the 2003-2004 academic year with significant progress in the first three areas: • Completion of of debuggingfortesting and hybrid the843Figure 9. Benchmarking results for hybrid fabrication (cid:150) casting of metal with fine holes.fabrication process shown in Fig. 7.Removal of the expendable mold material after casting without damaging the casting. • Reduction of risk of the expendable mold cracking or failing during metal pressurization. • Begin the evaluation of mechanical properties as a function of lattice element feature size and comparison of mechanical properties to wrought properties.4.

• Begin the evaluation of mechanical properties as a function of lattice element feature size and comparison of mechanical properties to wrought properties.4. CONCLUSIONSThe progress made during the first year and a half of work on this project suggests that this research will lead to practical procedures for the design and fabrication of lightweight, high performance structure and machine parts. The key to these new procedures is the further development of hybrid fabrication techniques to allow production of components with complex optimized microstructures. Significant hurdles remain (cid:150) in both the design and fabrication halves of the project.5. ACKNOWLEDGMENTSThe authors are pleased to recognize that the progress described in this paper demonstrates the outstanding contributions of Undergraduate Research Assistants Mark Batdorff, Adam Schneider, and Josh Rocholl.

In this research, selective laser melting (SLM) technology was used to fabricate structures with gradient lattice designs. The structure was varied in strut thickness continuously and linearly in single direction for cubic and honeycomb unit cells. Results showed that the complex design was successfully built and achieved nearly full- dense strut. Compression test results showed that the stress-strain curves of both cubic and honeycomb lattice structures oscillate with multiple peak loads, suggesting ductile characteristics. However, lattice structures with graded thickness tend to oscillate upward as the strut diameter increases.IntroductionRecently, lightweight metallic lattice structures have gained more attention due to their unique shock energy absorbing ability. This makes them suitable for use in applications like bone-mimicking biomaterials, heat exchanger or other load-bearing applications [1-4].

This makes them suitable for use in applications like bone-mimicking biomaterials, heat exchanger or other load-bearing applications [1-4]. However, the fabrication of metallic lattice structures is still challenging which requires novel manufacturing techniques to overcome inhomogeneity within the cell architectures. In addition, metallic lattice structures have very different deformation behavior from conventional dense metals, thus requiring a special set of compression testing methods for porous and lattice metals such as ISO standard 13314:2011 (Mechanical testing of metals -- Ductility testing -- Compression test for porous and lattice metals).Selective Laser Melting (SLM) is a layer additive manufacturing process that can be used to manufacture complex metal components. This is achieved by selectively melting metal powder, layer by layer, according to the 3D design model.

This is achieved by selectively melting metal powder, layer by layer, according to the 3D design model. One of the advantages of this manufacturing process is its ability to manufacture dense, complex metal parts in net shape, such as metallic lattice structures [1, 2, 5-8].A wide range of lattice structure designs exist, with widely varying shapes and internal geometries. The lightweight porous structures currently used at present can be broadly divided into two categories: first is the more conventional block structure which consists of straight vertical solid walls with a few holes [9]. The solid walls can be arranged to form rectangular, triangular, hexagonal or other shapes.

The solid walls can be arranged to form rectangular, triangular, hexagonal or other shapes. Secondly, lattice designs have a more open structure which allow for easier removal of metal powders, and usually allow for lattice structures with lower volume fractions to be built.Much work has been done in trying to optimize the function of metallic lattice structures, by varying the geometry, volume fraction, and unit cell size of the structure [5, 8]. This study intends to expand on previous works by proposing a new approach to optimizing metallic lattice structures, through the novel application of functionally-graded thickness along the vertical direction.

Varying the thickness of the walls of the structure2387would allow for opportunities to optimize the shock absorbing capabilities of the lattice structure while minimizing the amount of material used to build the high-performance functionally graded components.Experimental procedureFour groups of lattice structure were designed (C1, C2, H1, H2): two different orientations of cubic unit repeating structure and honeycomb unit repeating structure (Figure 1). For groups of C1 and H1, the struts were either parallel or perpendicular to the build direction (BD), which is along the z-axis. For groups of C2 and H2, lattice structures also consisted of angled struts. The size of cubic unit was 2 mm x 2mm x 2mm. The size of honeycomb unit was 1 mm for every strut of hexagonal face and 2 mm height.

The size of cubic unit was 2 mm x 2mm x 2mm. The size of honeycomb unit was 1 mm for every strut of hexagonal face and 2 mm height. Each group consists of the following samples: functionally graded components with strut diameter changing linearly from 0.4 mm to 1.2 mm, and structure with uniform strut diameter of comparable material volume as FGC. Average size of samples was 14.2 mm x 13.0 mm x 11.6 mm (length x width x height). Three duplicates were printed for each design.CubicC1 (FGC)BDC2 (FGC)BDHoneycombH1 (FGC)BDH2 (FGC)BDFigure 1 Designs of functionally graded components. Build direction is denoted as BD.The lattice structures were designed with 3-matics software (Materialise NV) and then processed in Magics software (Materialise NV) before printing. The processed files were sent to SLM 250 HL (SLM Solutions Group AG) for printing with Ti6Al4V powder (TLS Technik GmbH & Co.) as material.

The processed files were sent to SLM 250 HL (SLM Solutions Group AG) for printing with Ti6Al4V powder (TLS Technik GmbH & Co.) as material. The titanium alloy powder used was grade 23 with spherical particle size of 20 - 63 μm. Printing process parameters were 150 W laser power, 400 mm/s laser speed, 80 μm hatch spacing, and 30 μm layer thickness. After printing, the structures were separated from base plate by wire cutting.Density of lattice strut was determined by Archimedes’ principle using a weighing device (Mettler Toledo XS204) for measuring mass in air and mass in ethanol. Density of lattice structure was calculated by dividing mass in air with total volume of lattice structure obtained from multiplication of length, width and height of the structure. The calculated density of lattice structure was divided by 4.43 g/cm3 (theoretical density of bulk Ti6Al4V) to get the values of relative density.

The calculated density of lattice structure was divided by 4.43 g/cm3 (theoretical density of bulk Ti6Al4V) to get the values of relative density. Porosity was calculated by subtracting volume of solid portion from total volume of lattice structure. Volume of solid portion of the lattice structure was obtained by back calculating from the density of lattice strut and mass of lattice structure in air.Struts of lattice structure were observed by light optical microscope (Olympus Upright Microscope MX51). Compression tests were performed with universal mechanical testing machine (Instron 5500R, Instron 23884505, Shimadzu Autograph AG-X Plus) under 0.05 per minute strain rate. Some samples were cut to have smaller cross head area if machine force limit is exceeded.

Some samples were cut to have smaller cross head area if machine force limit is exceeded. Video recording or photographs taken at regular intervals were done during the compression test to identify failure behavior of the structures.Results and DiscussionThe metallic lattice structures of cubic and hexagonal cell geometries with functionally graded designs built by SLM were shown in Figure 2. Specimens with uniform strut diameters of comparable volume fractions were also built for comparison in compressive properties.BDBDBDBDC1 (FGC)C2 (FGC)H1 (FGC)H2 (FGC)Figure 2 Fabricated functionally graded components for compression testing. Build direction is denoted as BD.Table 1 shows the differences in porosity and density of the fabricated lattice structures versus the original design values.

Build direction is denoted as BD.Table 1 shows the differences in porosity and density of the fabricated lattice structures versus the original design values. As shown in the table, the design strut diameters of the lattice structures with uniform strut were 0.995 mm, 0.875 mm, 0.844 mm, 0.858 mm for C1 (uniform), C2 (uniform), H1 (uniform), H2 (uniform), respectively. However, the measured strut diameters were 0.97 mm, 1.02 mm, 0.86 mm, 0.98 mm for C1 (uniform), C2 (uniform), H1 (uniform), H2 (uniform), respectively. This is why the design porosity was higher than the average porosity of the fabricated part in most cases. It was observed some partially-melted metal powders bonded onto the struts, rather than forming solid and continuous strut. The irregularities caused by powder adhesion may affect the mechanical properties of the part fabricated by SLM [1, 8]. Nevertheless, the calculated density of lattice structure was typically higher than 99%.

Nevertheless, the calculated density of lattice structure was typically higher than 99%. Powder adhesion on the strut for cubic and honeycomb lattice structures is shown in Figure 3.

Due to the manufacturing constraints imposed by SLM, it should be noted that struts of lattice structures cannot be longer than 2mm without support or orient at an angle less than ~40° with respect to the base plate.Table 1 Porosity and density of the fabricated lattice structures versus design valuesDesignSampleStrut diameter (mm) MeasuredPorosity (%) Average porosity of part 65.1 ± 0.3 56.9 ± 0.2 64.2 ± 0.4 59.5 ± 1.8 44.9 ± 0.9 41.3 ± 0.4 39.0 ± 0.2 37.0 ± 0.4Porosity difference* 9.0 19.9 7.6 13.3 11.9 14.8 20.9 21.4 * Porosity difference = (Design porosity - Average porosity of part) / Design porosityDesignDesign porosity 71.5 71.0 69.5 68.7 50.9 48.4 49.3 47.1FGC Uniform FGC Uniform FGC Uniform FGC Uniform0.4 – 1.2 0.995 0.4 – 1.2 0.875 0.4 – 1.2 0.844 0.4 – 1.2 0.8580.968 ± 0.004 - 1.021 ± 0.022 - 0.858 ± 0.010 - 0.976 ± 0.036C1C2H1H22389Figure 3 OM images showing powder adhesions on the struts, resulting in rough surface finish.Figure 4-7 shows compression test results for different lattice structure designs, including cubic and honeycomb unit cells, different orientation, and various strut diameters – in graded and uniform shapes.

Results reveal that when there were more vertical struts compressed parallel to the build direction, the overall first maximum compressive strength were higher, i.e. C1 > C2, H1 > H2. This means lattice structures with larger amount of vertical struts parallel to the build direction were stiffer and sturdier if compressed along the build direction. Similar phenomenon was reported by McKown et al. as significantly higher modulus occurred due to the presence of the vertical pillars or struts [2].

The same trend was observed for both FGC and uniform lattice structures.C1 FGCC1 Uniform600600500sample 1 sample 2 sample 3500sample 1 sample 2 sample 3) a P M400) a P M400(s s e r t S300200(s s e r t S3002001001000002060 40 Strain (%)8010002060 40 Strain (%)80Figure 4 Compression test results for cubic lattice structures of design 1, denoted as C1.2390100) a P M(s s e r t S) a P M(s s e r t S3002502001501005001400120010008006004002000C2 FGCC2 Uniform300sample 1 sample 2 sample 3250sample 1 sample 2 sample 3) a P M200(s s e r t S15010050002040 60 Strain (%)8010002040 60 Strain (%)Figure 5 Compression test results for cubic lattice structures of design 2, denoted as C2.H1 FGCH1 Uniformsample 1 sample 2 sample 314001200sample 1 sample 2 sample 3) a P M(1000800s s e r t S600400200002040 60 Strain (%)8010002040 60 Strain (%)Figure 6 Compression test results for honeycomb lattice structures of design 1, denoted as H1.23918080100100800H2 FGC800H2 Uniform) a P M700600500sample 1 sample 2 sample 3) a P M700600500sample 1 sample 2 sample 3(s s e r t S400300(s s e r t S4003002002001001000002040 60 Strain (%)8010002040 60 Strain (%)80Figure 7 Compression test results for honeycomb lattice structures of design 2, denoted as H2.The results of the compression study clearly show the difference in stress-strain response between different lattice structure with different types of unit cell designs as well as the effects of load direction on different build directions.

In the cases of lattice structures with uniform strut diameters, typical stress-strain response that includes initial quasi-elastic gradient, plateau regions, as well as fluctuations of the stress-strain curve were observed. Similar, the stress-strain curve of the lattice structures with graded diameters is also composed of initial linear response, a number of plateaus and oscillations. However, the fluctuations were trending upward due to the increase in strut diameters, which translates to stronger regions that were more resilient to deformation. The results demonstrated that functionally graded components could potentially possess better shock absorbing ability as they could adeptly change the compressive response along the load direction.ConclusionsPreliminary results of design and fabrication of metallic lattice structures with cubic and honeycomb unit cells were presented.

It was observed that the fabricated part porosity was less than the porosity of design porosity, which was a result of partially-melted powder adhesion on the struts. Compression tests were also conducted on lattice structures with uniform strut diameters and graded strut diameters. The result showed a characteristic stress-strain response of the porous lattice structure. The influence of graded strut diameters was demonstrated as the fluctuations were trending upward, showing stiffer and stronger properties of the functionally graded structures. It is expected FGC can provide adaptable material properties in different loading conditions, as well as improved energy absorption capability while minimizing material usage.AcknowledgmentsThis work was financially supported by A*STAR Industrial Additive Manufacturing Program: WorkPackage 2 (Selective Laser Melting), under Grant No.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesign and Fabrication of Functionally Graded Material from Ti to γ-TiAl by Laser Metal Deposition Xueyang Chen1, Lei Yan1, Joe Newkirk2, and Frank Liou11Department of Mechanical and Aerospace Engineering, 2Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractFunctionally graded material (FGM) is one kind of advanced material characterized by a gradual change in properties as the position varies. The spatial variation of compositional and microstructure over volume is aimed to control corresponding functional properties. In this research, when 100% γ-TiAl was directly deposited on pure Ti substrate, cracks were formed within the γ-TiAl layer.

In this research, when 100% γ-TiAl was directly deposited on pure Ti substrate, cracks were formed within the γ-TiAl layer. Then a six-layer crack-free functionally graded material of Ti/TiAl was designed and fabricated by laser metal deposition (LMD) method, with composition changing from pure Ti on one side to 100% γ-TiAl on the other side. The fabricated FGM was characterized for material properties by a variety of techniques. The chemical compositions, microstructure, phases, and hardness of the composite were characterized by Scanning Electronic Microscope (SEM), Optical Microscope (OM), Energy Dispersive X-ray Spectroscopy (EDS), and hardness testing. The microstructure and chemical compositions in different layers were studied.1. IntroductionThe origin of functionally graded material (FGM) is from the National Aerospace Laboratory of Japan in 1984 during a spaceplane project.

IntroductionThe origin of functionally graded material (FGM) is from the National Aerospace Laboratory of Japan in 1984 during a spaceplane project. The proposed thermal barrier material is capable of withstanding a surface temperature of 2000 K and a temperature gradient of 1000 K across a 10 mm section [1]. FGM is one kind of advanced material characterized by a gradual change in properties as the position varies. The spatial variation of compositional and microstructure over volume is aimed at controlling corresponding functional properties, such as mechanical properties, thermal properties, electrical properties, etc. In FGM research, various methods to prepare and fabricate FGM materials are available. The fabrication methods include powder metallurgy [2], [3], thermal spraying [4] [5], additive manufacturing [6], [7]. Additive manufacturing is a manufacturing method that can create 3D objects.

The fabrication methods include powder metallurgy [2], [3], thermal spraying [4] [5], additive manufacturing [6], [7]. Additive manufacturing is a manufacturing method that can create 3D objects. This method allows you to produce complex products with minimal lead times directly from the 3D CAD models by adding layers of materials on top of each other. The bind layers schematic could mix elemental powders of variable ratios in separate layers to fabricate FGM parts with controllable gradients, no matter the complexity [8].There is a wide range of applications for FGM and different types of FGMs could be used in different areas of application. In aerospace and automotive area, the space shuttle utilized the FGM material made of ceramic and metal that can provide the thermal protection and load carrying capability to replace the ceramic tiles that easily to crack [9].

FGMs are also used as dental implants148in the biomedical area [10], vehicle armor in defense area [11], high-density magnetic recording media in optical applications [9], cutting tools in the industry [12] etc.Actually, compressors and turbine disks are two of the most significant parts in the aerospace industry. Turbine disks require different material properties at different regions. The rim region suffers from high temperature, requires good oxidation resistance. The core region has a relatively low temperature but requires high strength. The rim and the core have different requirements. The current approach is to use different treated nickel-based superalloys for each part. Nickel-based alloys are widely used in jet engines because it could withstand high- temperatures. In aerospace applications, Ti has a combination of excellent properties but has limitation for oxidation resistance.

In aerospace applications, Ti has a combination of excellent properties but has limitation for oxidation resistance. While, intermetallic titanium alloys such as γ-TiAl, Ti3Al can have higher oxidation resistance and both of the two alloys can be used under high temperature. Because the density for the γ-TiAl is much lower than the nickel-based superalloys and its specific properties, γ-TiAl tends to replace them. In this way, the rim can be made of Ti3Al or γ-TiAl, which have good high-temperature properties. And the core can be made of Ti with good strength at relatively low temperatures. However, there are some problems to join them effectively. Thus, the aim of this project is to manufacture Ti/γ-TiAl functionally gradient material (FGM) using Laser Metal Deposition (LMD) method. LMD is one of the advanced additive manufacturing technologies. It could fabricate complex near net shape metallic parts layer by layer directly from computer-aided design (CAD) models.

LMD is one of the advanced additive manufacturing technologies. It could fabricate complex near net shape metallic parts layer by layer directly from computer-aided design (CAD) models. This technique could reduce cost by reducing the buy-to- fly ratio and lead time of production.The main goals of this work involved studying the microstructure and mechanical properties of the fabricated alloys. In this study, commercially purified (CP) titanium plate was used as the substrate material for all the depositions. The powder material was fabricated as thin wall depositions and characterized for material properties using optical and Scanning Electron Microscope (SEM) microstructure characterization, hardness testing, Energy Dispersive X-ray Spectroscopy (EDS). The study investigated the material properties of γ-TiAl manufactured by LMD method. And it showed the possibility to control properties and crack formation in the deposits.

The study investigated the material properties of γ-TiAl manufactured by LMD method. And it showed the possibility to control properties and crack formation in the deposits. Also, it demonstrated the approach to build a complex FGM system (Ti + γ-TiAl and/or any intermediate FGM materials in between) by LMD method with hope to move to more exotic material compositions in the future.2. Materials and Methods2.1 Material PreparationUnder the current study, commercially purified gas atomized γ-TiAl (Ti-48Al-2Cr-2Nb) powder and pure Ti powder were used as precursor materials. For this kind of γ-TiAl powder, the addition of Cr element can increase the ductility of the material. The addition of Nb element is to enhance the oxidation resistance. Blends of the two kinds of powder with different composition were made to achieve layers in functionally graded materials (FGM). The particle size distributions of γ-TiAl and Ti powder are listed in Table 1 and Table 2 individually.

The particle size distributions of γ-TiAl and Ti powder are listed in Table 1 and Table 2 individually. The chemical compositions of the two kinds of powder are given in Table 3 and Table 4. Commercially purified (CP) titanium plate was used as the substrate material for these depositions, and substrates were prepared to the dimensions of 2 × 0.5 × 0.25 in.149Table 1. Particle size distribution of Ti-48Al-2Cr-2Nb powder (% Under).Particle Size (µm) 150 10675 99.9 66.2 25.5 14.7 4.06353 Apparent Density (g/cc)By Mass2.20Table 2. Particle size distribution of Ti powder (% Under).Particle Size (µm) 150 53 Apparent Density (g/cc)By Mass98.9 0.92.64Table 3. Chemical analysis of Ti-48Al-2Cr-2Nb powder (wt.%).Elements Measured Balance 34.4 2.38 4.75 0.04 0.016 0.014 0.128TiAl Cr NbFeSiCOTable 4.

Chemical analysis of Ti-48Al-2Cr-2Nb powder (wt.%).Elements Measured Balance 34.4 2.38 4.75 0.04 0.016 0.014 0.128TiAl Cr NbFeSiCOTable 4. Chemical analysis of Ti powder (wt.%).Elements Fe Other Measured Balance <0.01 <0.16 <0.01 <0.001 <0.12 <0.4TiCONH2.2 Deposition SystemIn this study, a 1 kW continuous wave fiber laser with a wavelength of 1064 nm was used to deposit the elemental powder blends. A melt pool spot size of approximately 2 mm was obtained through a lens of 750 mm focal length. A ceramic tube nozzle was used in conjuncture with a CNC (Computer Numerical Control) table to facilitate movement and perform deposition. Argon gas was used to implement an inert atmosphere and was also used as a carrier gas to deliver the powder mixture to the melt pool.The schematic layout of the LMD system is shown in Figure 1. LMD deposition is an additive manufacturing technique capable of fabricating complicated structures with superior properties [13].

LMD deposition is an additive manufacturing technique capable of fabricating complicated structures with superior properties [13]. The applications of this technique include coatings, rapid prototyping, tooling, repair, etc. LMD uses a focused laser beam as a heat source to create a melt pool into which powder feedstock is injected. The powder material is metallurgically bonded to the substrate through solidification [14].Figure 1. Schematic layout of the LMD (Laser Metal Deposition) system [15].1502.2 Experimental ProcedureIn this research, three kinds of experiments have been designed in order to get crack-freeTi/γ-TiAl deposit.In the first kind of experiment, set-up 1 has been used to deposit Ti-48Al-2Cr-2Nb powder on CP Ti substrate directly, as Figure 2 (a) showed below. In this set-up, the substrates were directly connected with the fixture. Thin wall depositions were fabricated by performing single melt pool layer by layer.

In this set-up, the substrates were directly connected with the fixture. Thin wall depositions were fabricated by performing single melt pool layer by layer. Before deposition, a preheat scan of running the laser across the substrate surface at 1 kW power without powder was performed, in order to ensure perfect bonding. The preheating process was also used to heat up the substrate and rid the surface of the substrate from oxide scale buildup, surface impurities, and, surface imperfections. Laser transverse speed was set at 600 mm/min for all the depositions. The laser power was set at 600 W. The initial 2 layers of deposition carried out a duty cycle of 100% to ensure perfect bonding with the substrate. The next 4 layers keep the duty cycle of 80% and then drop to 60% for another 4 layers. The rest of the deposition was carried at a duty cycle of 40%. The actual laser power is calculated by multiplying the laser power by the duty cycle.

The rest of the deposition was carried at a duty cycle of 40%. The actual laser power is calculated by multiplying the laser power by the duty cycle. The same deposition scheme and parameters were used for all the depositions. This was expected to ensure an unbiased comparison between different depositions. The second kind of experimental design used the set-up 2 as shown in Figure 2 (b). The insulating refractory bricks were inserted between the substrate and the fixture.(a) Set-up 1(b)Set-up 2Figure 2. Two experimental set-upsThe third experimental design was used to manufacture the crack-free functionally gradient material (FGM) from Ti to γ-TiAl. The FGM study included material design, material preparation, and evaluation of material properties. The FGM path was developed as shown in Figure 3. The atomic percentage of each element and possible phase composition of every layer in the FGM deposit are also listed in Table 5.

The FGM path was developed as shown in Figure 3. The atomic percentage of each element and possible phase composition of every layer in the FGM deposit are also listed in Table 5. Powders were weighed for desired compositions in a glove box under an argon atmosphere and sealed in bottles before starting the deposition process. In this study, γ-TiAl powder and pure Ti powder were used. The as-blended powder mixtures can be used151to form the layers in the functionally graded materials. These bottles were then shaken using a Turbula® mixer (Glen Mills Inc., Maywood, NJ, United States) for 20 min to obtain thorough mixing and homogeneity within the powder blends. During the process, there is no insulating brick used here by taking use of the first experimental set-up (set-up 1).Figure 3. The designed functionally graded material (FGM) pathTable 5.

During the process, there is no insulating brick used here by taking use of the first experimental set-up (set-up 1).Figure 3. The designed functionally graded material (FGM) pathTable 5. Particle size distribution of Ti-48Al-2Cr-2Nb powder (% Under).Weight Percentage 100% γ-TiAl 10% Ti + 90% γ-TiAl 30% Ti + 70% γ-TiAl 50% Ti + 50% γ-TiAl 70% Ti + 30% γ-TiAl 90% Ti + 10% γ-TiAlTi at.% Al at.% Cr/Nb at.% Possible Phase Composition48.00 52.30 61.40 71.28 82.01 93.7348.00 44.03 35.63 26.51 16.60 5.792.00 1.83 1.48 1.10 0.69 0.24α2 + γ α2 + γ α2 + γ α2 α2 + α αAfter deposition, vertical transverse sections were cut using a wire Electro-Discharge Machine (wire-EDM, Hansvedt Industries Inc., Rantoul, IL, United States) and mounted in Bakelite for grinding and polishing. Kroll’s reagent containing 30 ml H20, 10ml HNO3, and 10 ml HF was used for etching the samples. Hirox optical microscope was used to perform microstructure imaging.

Kroll’s reagent containing 30 ml H20, 10ml HNO3, and 10 ml HF was used for etching the samples. Hirox optical microscope was used to perform microstructure imaging. Helios Nanolab 600 SEM (FEI Company, Hillsboro, OR, United States) equipped with an Oxford Energy Dispersive Spectrometer was used to get the chemical compositions. The Vickers hardness measurements were performed using a Struers Duramin micro-hardness tester (Struers Inc., Cleveland, OH, United States). Press load of 9.81 N and load time of 10 s were used during the hardness tests.3. Results and discussion3.1 Deposit ProfileFigure 4 shows the deposited samples in the first experimental design. The image of the γ- TiAl deposit made by using set-up 1 in the first design has been shown in Figure 4 (a). It could be seen that there are some cracks in the deposit. Then with the same setup, another deposited sample was made from the powder blend of 80 wt.% γ-TiAl +20 wt.% Ti.

It could be seen that there are some cracks in the deposit. Then with the same setup, another deposited sample was made from the powder blend of 80 wt.% γ-TiAl +20 wt.% Ti. Figure 4 (b) demonstrates the 80 wt.% γ-TiAl +20 wt.% Ti sample deposited by using set-up 1. It has been found that it is possible to make crack-free deposits with the blend. It is theorized that the cracks that mainly exist152in the vertical direction are caused by residual stress. It is known that there exists a great difference in the thermal expansion coefficient between Ti and γ-TiAl. Ti has the thermal expansion coefficient of 8.9 × 10−6𝐾−1, while γ-TiAl shows 14.4 × 10−6𝐾−1. Due to the high residual thermal stresses between the Ti and γ-TiAl layers, it would result cracks during the cooling process after deposition. However, the layer of Ti + γ-TiAl deposits could play a role in reducing the interfacial residual thermal stresses.

However, the layer of Ti + γ-TiAl deposits could play a role in reducing the interfacial residual thermal stresses. Therefore it could cause some cracks in the deposit by directly depositing γ-TiAl powder on Ti substrate, however crack-free γ-TiAl deposit could be made on the Ti substrate by using the powder blend of γ-TiAl and 20 wt.% Ti.(a) γ-TiAl sample without insulating bricks(b) 80 wt.% γ-TiAl +20 wt.% Ti sample without insulating bricksFigure 4. Deposited samples in the first experimental designThen in order to get a crack-free γ-TiAl deposit, the set-up 2 has been used in the second experimental design, as shown in Figure 2 (b). The image of the γ-TiAl deposit made by using set- up 2 has been shown in Figure 5. It has been found that the deposited sample has no cracks. In this set-up, the insulating bricks were inserted between the substrate and the fixture.

It has been found that the deposited sample has no cracks. In this set-up, the insulating bricks were inserted between the substrate and the fixture. Because these materials thermally isolate the substrate, the bricks could retain more of the input laser energy and decrease cooling rate. Then many conditions, such as boundary conditions, thermal insulation, and post-process heat treatment, may affect the mechanical properties of deposits.Figure 5. Deposited sample in the second experimental designHowever, it is not very flexible to fabricate some big and complex parts with insulating bricks in the industry area. It is necessary to develop a gradient path to realize the same target of depositing crack-free γ-TiAl deposit on Ti substrate. Then an FGM path has been designed, as shown in Figure 3.

It is necessary to develop a gradient path to realize the same target of depositing crack-free γ-TiAl deposit on Ti substrate. Then an FGM path has been designed, as shown in Figure 3. The objective of the designed FGM is to meet the expected requirements by eliminating the obvious interface between composite materials in different layers, just the same as bones, teeth, etc. in nature. Also, the FGM is also designed to accommodate materials for specific153applications. During the FGM deposition process, there was no insulating brick used. From Figure 3, the FGM parts were made with five different powder blends and the top layer is 100% γ-TiAl. Along the deposit, there are six regions with different chemical composition. The gradients between different layers could smooth transition to reduce the opportunity of failure. The transient layer of Ti + γ-TiAl deposits could gradually reduce the residual thermal stresses.

The gradients between different layers could smooth transition to reduce the opportunity of failure. The transient layer of Ti + γ-TiAl deposits could gradually reduce the residual thermal stresses. Figure 6 showed the image of the FGM deposit and its transverse section after cutting.(a) FGM sample(b) Transverse section of the FGM depositFigure 6. Deposited sample in the FGM design3.2 Vickers hardnessFigure 7 shows the Vickers hardness distribution of the deposited samples without using insulating bricks in the first experimental design. By applying post heat treatment during the deposition process, the Vickers hardness values of the deposits could also be compared in this Figure. It was observed that the Vickers hardness values of the γ-TiAl layers varied around 375 HV in the middle part of the deposit zone, then increased to 450 HV as it progressed into the top of the deposit.

While the deposit by using the powder blends of 80 wt.% γ-TiAl +20 wt.% Ti had a Vickers hardness of around 420 HV. The literature reported γ-TiAl Vickers hardness was 425 HV. And when comparing the two kinds of deposits under different conditions of performing post heat treatment or not, it has been found that the post heat treatment on deposit would increase the hardness values, but could not make significant changes.154Figure 7. Vickers hardness distribution of the deposited samples in the first experimental design.Figure 8 shows the hardness distribution of the deposited samples under different laser power with using insulating bricks in the second experimental design. It can be observed that when other parameters stayed the same, the higher the laser power, the lower the Vickers hardness. The reason should be the increase of the energy input. It would decrease the cooling rate, and then cause coarser microstructure and get lower hardness values.

The reason should be the increase of the energy input. It would decrease the cooling rate, and then cause coarser microstructure and get lower hardness values. Therefore with insulating bricks to decrease the cooling rate, it is much easier to get crack-free γ-TiAl deposits.155Figure 8. Vickers hardness distribution of the deposited samples in the second experimental design.Functionally graded materials (FGM) is one kind of advanced material. They have a gradual change in properties as the position varies. As shown in this Figure 3, the volume fraction of one material will change from 100% on one side to zero on another side, and another constituent will change in the other way around. Their composition and microstructure vary in space by following a predetermined law. This change also leads to a gradient of properties and performance, such as mechanical properties, thermal properties. Figure 9 shows the Vickers hardness distribution of the FGM sample.

This change also leads to a gradient of properties and performance, such as mechanical properties, thermal properties. Figure 9 shows the Vickers hardness distribution of the FGM sample. The 100% TiAl layer is harder than previous composition layers.156Figure 9. Vickers hardness distribution of the FGM sample3.3 Microstructure & EDS testThe microstructures of γ-TiAl deposits from Set-up 1 and Set-up 2 are shown in Figure 10.It shows that most prevalent microstructure for both is lamellar.(a) Microstructure of the γ-TiAl from set-up 1(b) Microstructure of the γ-TiAl from set-up 2Figure 10. Microstructures of the γ-TiAl from two set-upsAnd Figure 11 shows the EDS result of the FGM deposit. Line scan was performed along the deposit, it differentiated six regions with different compositions listed in Table 5. The designed length for each region is 5 mm. Region height difference can be attributed to the powder capture efficiency.157Figure 11.

The designed length for each region is 5 mm. Region height difference can be attributed to the powder capture efficiency.157Figure 11. EDS result of the FGM deposit4. ConclusionsThe current research was implemented to investigate the feasibility of combining Ti with γ-TiAl via the LMD process. The fabricated material was characterized using techniques such as scanning electron microscopy, hardness testing, EDS. The conclusions from the analyses include:Crack-free γ-TiAl deposits could be fabricated by using insulating bricks to control the cooling rate during LMD process.Proposed an FGM path to successfully join titanium alloy with γ-TiAl. • Mechanical properties of γ-TiAl would be affected by are affected by processing parameter. Increasing energy input would decrease the cooling rate and decrease the Vickers hardness.The microstructure in the γ-TiAl deposit is lamella.5.

SIGN AND FREEFORM FABRICATION OF DEPLOYABLE STRUCTURES WITH LATTICE SKINS Uma Maheshwaraa, Catherine Tradd, David Bourell, and Carolyn Conner Seepersad Mechanical Engineering Department The University of Texas at Austin, Austin, TX 78712AbstractFrontier environments—such as battlefields, hostile territories, remote locations, or outer space—drive the need for lightweight, deployable structures that can be stored in a compact configuration and deployed quickly and easily in the field. We introduce the concept of lattice skins to enable the design, solid freeform fabrication (SFF), and deployment of customizable structures with nearly arbitrary surface profile and lightweight multi-functionality. Using Duraform FLEX® material in a selective laser sintering machine, large deployable structures are fabricated in a nominal build chamber by either virtually collapsing them into a condensed form or decomposing them into smaller parts.

Before fabrication, lattice sub-skins are added strategically beneath the surface of the part. The lattices provide elastic energy for folding and deploying the structure or constrain expansion upon application of internal air pressure. Nearly arbitrary surface profiles are achievable and internal space is preserved for subsequent usage. In this paper, we present the results of a set of experimental and computational models that are designed to provide proof of concept for lattice skins as a deployment mechanism in SFF and to demonstrate the effect of lattice structure on deployed shape.1. IntroductionSeveral mechanisms are available for compactly storing and deploying two-dimensional surfaces and three-dimensional structures for applications ranging from the deployment of satellite booms and solar arrays to the construction of temporary shelters and retractable stadium roofs with vast unsupported spans [1,2].

Examples include tensegrity structures comprised of a combination of exclusively compressive and tensile members, as observed in structures such as the Georgia Dome in Atlanta, and pantographs—structural mechanisms with bars connected by pivots at the center point and ends of each bar, and embodied in collapsing laundry racks, temporary gates, and the popular children’s toy, the Hoberman sphere [1,2].Conventional deployment mechanisms and fabrication techniques generally restrict the geometry of a deployed structure to symmetric, polygonal or spherical shapes and make it difficult to rapidly customize the geometry and functionality of the device.

In contrast, a freeform fabrication approach offers several potential advantages, including: (1) the capability of fabricating complex, intricate internal and external geometries, enabling a host of customized, embedded deployment mechanisms and customized parts with nearly arbitrary deployed lightweight, geometries, (2) rapid design and fabrication cycles, and (3) potentially multifunctional structures with embedded mechanisms that not only control deployment but also provide additional functionality such as strength, stiffness, insulation, or blast amelioration.

Freeform deployment also has the potential to overcome build chamber size limitations by building large parts in a condensed configuration and deploying them, post-build.598We seek to establish “Design for Freeform Deployment” methods and accompanying solid freeform fabrication (SFF) techniques to enable the design, fabrication, and deployment of customizable structures with nearly arbitrary surface geometry and lightweight multi- functionality. The first step towards this goal is to identify potential deployment mechanisms that are suitable for freeform fabrication and support the desired capabilities. For this purpose, we have devised a particularly promising concept that uses lattice sub-skins for deployment, as illustrated in Figure 1. Beginning with a part of arbitrary surface profile and hollow interior, lattice sub-skins are added beneath the surface of the structure.

Beginning with a part of arbitrary surface profile and hollow interior, lattice sub-skins are added beneath the surface of the structure. The part is fabricated using selective laser sintering (SLS) technology and a flexible, elastomer material called Duraform® FLEX. If the structure is larger than the build chamber, it is decomposed and fabricated as a collection of parts that are subsequently joined together. The flexible structure can be folded for ease of storage and transport and then deployed in the field via a combination of elastic strain energy and pneumatics. In this process, the lattice structure serves several functions. First, during the folding step, it stores strain energy that can be returned upon unfolding to help deploy the structure into its original configuration. Second, in its deployed form, the lattice structure supports the surface of the flexible part to prevent collapse and distortion of desired surfaces.

Second, in its deployed form, the lattice structure supports the surface of the flexible part to prevent collapse and distortion of desired surfaces. Finally, if elastic energy is insufficient for deploying a large structure under its own weight, air pressure is applied inside the structure, and the lattice skin constrains the expansion of the structure to prevent balloon-like inflation and preserve desired surface profiles. An alternative fabrication route is to collapse the structure virtually (much like deflating a balloon), fabricate it in its condensed form, and then use internal air pressure to inflate the part, with the lattice structure constraining its expansion and preserving the desired surface profile. Fabrication by decomposition is emphasized in this paper.Figure 1.

Fabrication by decomposition is emphasized in this paper.Figure 1. Airfoil section with lattice sub skinLattice structures and cellular materials are particularly amenable to solid freeform fabrication techniques, which provide flexibility for fabricating complex, customized lattices or cellular structures as integral aspects of an overall part. Lattice structures and cellular materials have been fabricated with a variety of SFF technologies, including SLS [3], direct metal deposition [4], selective laser melting [5], stereo/photolithography techniques [6,7], and hybrid approaches that include SFF-based sacrificial molds [8,3]. They are used for applications that require lightweight stiffness, compliance, impact absorption, heat exchange, catalysis, and other properties (cf. [9-11] for an overview).

They are used for applications that require lightweight stiffness, compliance, impact absorption, heat exchange, catalysis, and other properties (cf. [9-11] for an overview). However, the use of lattice structures as deployment mechanisms has not been investigated to date.Lattice structures have several characteristics that make them particularly appropriate for deployment applications. First, they are preferable to internal mechanism-like structures, mentioned previously, because they are more flexible for shape control of relatively arbitrary,599freeform geometries; occupy very little space in the build chamber; and do not require small- scale pivots or joints that are difficult to fabricate.

Second, they also offer a combination of low relative density and high effective stiffness, providing high levels of rigidity for controlling the deployment of a three-dimensional part without significantly impacting its apparent density for Third, the lattice skins do not occupy the interior regions of the consolidation and collapse [12]. part, leaving the space available for other uses. In addition, the lattice skin itself makes the deployed part suitable for multifunctional applications [9,11,13,14]. For example, the lattice sub-skin could be filled with cooling fluids or air for heat exchange or insulation; earth, foam, or other materials for impact absorption or blast protection; or explosive materials for controlled detonation.

Finally, lattice skins are conducive to portable deployment, requiring only a portable air pump for deployment and a spray can of thermoset coating for rigidization, if desired.In the following sections, a method is presented for designing, fabricating, and deployingstructures with internal lattice sub-skins, and preliminary results are presented.2. Method for Design and Freeform Fabrication of Deployable Structures with Lattice Sub-skinsArbitrary 3D Surface1. Add Lattice Elements2. Optimize Lattice Structure3a. Decompose into Small Parts Virtually3b. Collapse Structure Virtually (Optional)4. Selective Laser Sintering (SFF)5. Deploy6. StiffenFully Functional PartFigure 2. A method for designing and fabricating deployable structures.600The general method for designing and fabricating deployable structures with lattice sub-skins is outlined in Figure 2. The starting point for the method is the 3D surface profile of the part to be deployed.

The starting point for the method is the 3D surface profile of the part to be deployed. In Step 1, a lattice structure is added beneath the surface of the part as a deployment mechanism for deploying the part. The arrangement of the lattice structure could be determined with a topology optimization procedure or a standardized strategy could be applied. An example of a standardized strategy is illustrated in Figure 3, in which the lattice structure is composed of elements oriented perpendicular to the surface of the part and connected with another set of elements that are parallel to the surface of the part. Two types of lattice sub-skins are described in Section 3: (1) an open lattice arrangement, as illustrated in Figure 3, for directly reinforcing the surface of the part, and (2) a closed lattice arrangement for pneumatic inflation.Figure 3.

An arbitrary 3D part with added lattice structure.In Step 2, a finite element model of the part and lattice sub-skin is built to analyze the stiffness and deflection profile of the part. If adjustments are needed, the dimensions and density of the lattice structure are varied iteratively until the intended deflection is attained. This step can be executed manually or with an optimization algorithm. For the lattice structure illustrated in Figure 3, the stiffness of the lattice structure should be such that it supports the self- weight of the part with minimal deviation from the intended surface profile. If internal air pressure is used to deploy a closed part, the lattice structure should constrain expansion under internal air pressure so that the intended surface profile is achieved as closely as possible.The deployment strategy is intended to be applied to large parts with volumes that are greater than that of the SLS build chamber.

Therefore, Step 3 offers two options for overcoming this challenge. As illustrated in Figure 5, the first option is to decompose the structure virtually into parts that are smaller than the build chamber and then assemble and adhere the parts in a post- processing step and fold the flexible structure into a compact, storable package. The second option is to virtually collapse the part in CAD, prototype it in compressed state and expand in a post-processing step to get the final part. This option is similar to wadding or folding a piece of paper and then unfolding it after the build process with a combination of thermal and mechanical601loading. Figure 4 shows a virtually compressed flat plate post-processed to get a part bigger than build chamber.Figure 4. Virtually compressed accordion post processed into a flat plateDecomposed VirtuallyPrototypedPost ProcessedFigure 5.

Virtually compressed accordion post processed into a flat plateDecomposed VirtuallyPrototypedPost ProcessedFigure 5. Virtual decomposition of parts larger than the build chamber.After the part is virtually compressed or decomposed into smaller parts, it is fabricated in the SLS machine using Duraform Flex material in Step 4. This step also includes post-processing (e.g., assembling and adhering parts and spraying with an infiltrant to make the parts air-tight).602In Step 5, the prototyped part is deployed using internal air pressure and/or any form of mechanical assistance (e.g., human-assisted unfolding). Since the part is constructed of Duraform® FLEX, its skin is quite pliable—a useful feature for folding and storing the part but not for maintaining a rigid structure. If a stiffer skin is desired, the part’s skin can be coated with a thermoset polymer in Step 6. The outcome of the method is a fully functional, deployable part.3.

If a stiffer skin is desired, the part’s skin can be coated with a thermoset polymer in Step 6. The outcome of the method is a fully functional, deployable part.3. Lattice StructureAn integral part of the methodology for building deployable parts is the deployment mechanism. The lattice structure provides elastic stiffness for supporting the surface of the part and for constraining the expansion of a part during inflation. It can also store elastic strain energy as a large structure is folded and return it as the part is unfolded to aid in the deployment process. Two types of lattice structures are proposed, an open lattice structure (Figure 6a) and a closed lattice structure for air deployable parts (Figure 6b).(a)(b)Figure 6. Open lattice (a) for a self-supporting structure and closed lattice (b) for an air deployable part.3.1 Open Lattice StructureAn open lattice structure is composed of lattice elements added to the inner surface of a part.

The arrangement of elements may be determined with a topology optimization routine or by selecting a standard strategy. In the standard strategy illustrated in this paper, lattice elements emanate from the inner surface of a part in an orientation normal to the surface. Their ends are connected by similar lattice elements of same or different dimensions. Rows consisting of numerous lattice elements are added to span the entire surface of the part. Finally the mesh of lattice elements forming the lattice structure is obtained by connecting all the rows. Figure 7 shows the open lattice structure unit for a section of cylindrical surface.Lattice element normal to surfaceLattice interconnectionFRONT VIEWSIDE VIEWISOMETRIC VIEWFigure 7. Open lattice unit.603The lattice structure is added to the inner surface in an iterative procedure. The inner surface of a 3D model is used as a reference to create the lattice structure in SolidWorks 2005.

The inner surface of a 3D model is used as a reference to create the lattice structure in SolidWorks 2005. In order to maintain the shape of a 3D part, lattice structures are strategically placed beneath the surface to provide support to the surface after deployment. A single lattice element is created normal to the surface at a point. A linear array of the lattice element is added with the curve representing the cross section as a guide. This is then extended along the depth of the curve to generate lattices over the entire inner surface. The spacing of the elements in all directions is decided based on the lattice density required for the part.The dimensions and density of the lattice structure play a vital role in making the deployable part. The lattice element density beneath the surface depends not only on the geometry of the surface and the target deflection profile but also on the stiffness of the Duraform® FLEX material.

The lattice element density beneath the surface depends not only on the geometry of the surface and the target deflection profile but also on the stiffness of the Duraform® FLEX material. The laser power of the SLS machine can be adjusted to alter the stiffness of the Duraform® FLEX material. If the material is highly flexible, the lattice elements are more densely spaced and/or larger in dimension to provide more support to the flexible surface.Finite element analysis (ANSYS [15]) is used to determine the appropriate density and dimensions of the lattice structure for maintaining the surface profile of the final part. The procedure is iterative. The initial dimension of the lattice element is set to 2 mm x 2 mm because the resolution of the SLS machine can be poor for smaller elements. Depending upon the surface deflection estimates, the lattice element sizes and densities are iteratively adjusted until the deflections are acceptable.

Depending upon the surface deflection estimates, the lattice element sizes and densities are iteratively adjusted until the deflections are acceptable. Eventually, the iterative procedure will be guided by an optimization method that may provide better results faster. Figure 8 shows the iterative procedure to generate lattice structure for any part. The ANSYS models are described in Section 4.Import 3D model of partAdd Lattice ElementsInitial dimension of lattice 2 mm x 2mmDeflection of the model using ANSYSIncrease the size of lattice ElementsIs the model able to withstand its own weightNOYESPrototypeFigure 8.

Methodology for generating lattice structure.6043.2 Closed Lattice Structure for Pneumatic DeploymentIf air pressure is used to deploy the structure, a closed lattice structure is created to join the outer surface of the part with a concentric, inner surface that is separated from the outer surface everywhere by a set distance, as illustrated in Figure 9. For deployment, pressurized air is pumped between the concentric skins. Lattice structures join the surfaces at strategic points, and they are oriented to be perpendicular to both skins. The lattice structure mainly helps in maintaining the shape when air is pumped in between the parallel surfaces. Without the lattice structure, the skin would expand naturally to a sphere-like shape, with maximum ratio of volume to surface area. The lattice skin constrains the expansion so that the volume between the two surfaces is a maximized, which occurs when the skins are concentric and the lattice elements are stretched to maximum length.

Figure 9 shows a section of an air foil with lattice structure connecting inner and outer surface.Figure 9. Deployable structure with air as the deployment mechanism.The final part will have non-uniform densities of lattice elements, with the arrangement depending on curvature and desired surface accuracy. Broader spacing essentially provides less stiffness between the skins, allowing the outer skin to expand more when air is pumped into the part. Broader spacing can yield bulges in the surface of the part between lattice elements, similar to the bulges in a riveted pillow.

Broader spacing can yield bulges in the surface of the part between lattice elements, similar to the bulges in a riveted pillow. If the accuracy requirements or curvature rate is higher in a particular section, the lattice elements in that section are more densely arranged to provide tighter control over the surface profile during deployment.4.Proof of Concept with Virtual and Physical PrototypesTo evaluate the feasibility and effectiveness of the proposed deployment strategy, we developed finite element models of sample deployable structures in ANSYS and created corresponding prototypes with SLS and Duraform® FLEX material.4.1 Finite Element Modeling in ANSYSFinite element analysis of Duraform® FLEX parts needs to account for the elastomeric behavior of the material with its relatively large Poisson’s ratio and large displacements under relatively small applied loads.

To model the behavior of a part composed of Duraform® FLEX that undergoes large deformations, we developed finite element analysis models in ANSYS and utilized its geometrically nonlinear analysis capabilities. Before analyzing deployable structures, we calibrated our models with some simple experimental tests. A cantilever beam test was done on a Duraform® FLEX block of dimensions 93.5 mm x 32.5 mm x 3.5 mm. As shown in Figure 10, the boundary conditions included fixed deflection and rotation at one end of the specimen605and a set of increasing loads applied to the opposite end. Data for deflection as a function of applied load are recorded in Table 1. Material properties of the specimens were also evaluated and listed in Table 2. These material properties vary according to SLS build parameters such as laser power [16][17].

Material properties of the specimens were also evaluated and listed in Table 2. These material properties vary according to SLS build parameters such as laser power [16][17]. Accordingly, the properties were evaluated from the set of test parts, and the same build parameters were used for prototyping.93.5 mmLOAD3.5 mmDEFLECTIONXYFigure 10. Schematic diagram for cantilever beam test.Table 1. Displacement in y-direction.Load (N)0.16785 0.355231 0.462616 0.576859DEFLECTION (mm) 24.61 30.22 30.98 35.12Table 2. Material properties for ANSYS analysis.PropertyValueEX (Tensile Modulus) Density Poisson Ratio3.8 MPa 486.1 kg/m3 0.450.85223937.08A corresponding three-dimensional finite element model of the specimen was developed in ANSYS, using the appropriate material properties. Both large deformation and small displacement analyses were conducted, and the results were compared with experimental data in Figure 11.

Both large deformation and small displacement analyses were conducted, and the results were compared with experimental data in Figure 11. For both analyses, a SOLID45 element was used with element size of 0.15 mm (measured by edge length). In a small displacement analysis, the load is applied in a single step, and it is assumed that the geometry of the part is unchanged—an assumption that leads to very little error if the deformation is small. When the large deformation analysis is enabled in ANSYS, the load is applied incrementally in a series of steps. After each step, ANSYS updates the geometry to reflect the simulated deformations. The number of load steps for the large deformation analysis was set to 200.As shown in Figure 11, the large deformation model matches the experimental data much more closely than the small displacement model over a broader range of displacements.

The small displacement model is very accurate for small displacements, but increasingly overestimates deflection as the magnitude increases. As shown in Figure 11, the force-deflection curve for the small displacement model is linear, as opposed to the nonlinear relationship observed experimentally. The large deformation model is relatively accurate over the entire range of displacements and typically approximates the experimentally measured deflection606within 10%. With the low magnitude of Duraform® FLEX’s elastic modulus, large deflections are expected and may even occur under self-weight for large structures.The simple experiment confirms our capability of analyzing the force-deflection behavior of simple structures composed of Duraform® FLEX material, using large deformation analysis in ANSYS. The next step is to demonstrate how we can use this capability to analyze and design fully deployable structures.Figure 11.

The next step is to demonstrate how we can use this capability to analyze and design fully deployable structures.Figure 11. Comparison of deflection values between experiments and ANSYS.4.2 Physical PrototypesTwo different arch prototypes were used to study the ability of lattice structure to maintain the required profile. The first prototype—illustrated in Figure 12 and Figure 13—is a single arch with overall dimensions of 30 cm (length) x 8 cm (height) x 0.3 cm (thickness of skin). Prototypes of the arch were fabricated with and without the underlying lattice structure, with identical dimensions for the length, height, and thickness of the outer skin. Both prototypes were fabricated in an identical arched shape. As shown in Figure 12, the arch collapses under its own weight without the underlying lattice structure. The center of the lattice prototype deflects vertically by 2.1 cm under its own weight.

As shown in Figure 12, the arch collapses under its own weight without the underlying lattice structure. The center of the lattice prototype deflects vertically by 2.1 cm under its own weight. As shown in Figure 13, ANSYS models of the structure predict a deflection of 2.0 cm, within 5% of the actual experimental value. In the ANSYS model, the effect of self weight was simulated by adding gravitational body forces and specifying the measured density of the Duraform® FLEX material. Boundary conditions included constraints on horizontal and vertical deflection on one end of the arch and constraints on vertical deflection on the other end.The second prototype—illustrated in Figure 14-Figure 17—is a double arch with overall dimensions of 30 cm (length) x 7 cm (height) x 0.3 cm (thickness of skin). The purpose of constructing a second prototype was to investigate a more arbitrary surface profile with convexities and concavities in the surface.

The purpose of constructing a second prototype was to investigate a more arbitrary surface profile with convexities and concavities in the surface. As illustrated in the figures, prototypes of the double arch were also fabricated with and without the lattice structure, with all other features identical.607As shown in Figure 14 and Figure 17, the prototype without the lattice structure collapsed under its own weight to a single, concave arch. As shown in Figure 15 and Figure 16, the prototype with the lattice structure retained its shape. The center of the prototype deflects vertically by 1.1 cm under its own weight. ANSYS models of the structure predict a deflection of 1.1 cm, within 1 % of the actual experimental value. Boundary conditions were identical to those of the single arch model.A larger prototype of the single arch was also created with dimensions of 90 cm (length) x 30 cm (height) x 0.9 cm (thickness of skin).

The deflection in the centre of the prototype was measured to be 19.8 cm due to its own weight. ANSYS models of the prototype predict a deflection of 18.7 cm, within 5% of the actual experimental value. The boundary conditions were identical to the previous models.Figure 12. Arch prototype with and without lattice structure.Figure 13. ANSYS model of the arch withstanding its own weight. Colors represent displacement contours with blue contours on the order of 2 cm for an arch with a span of 30 cm and height of 8 cm.608Figure 14. ANSYS result for the double arch without lattice structure, showing deflection contours under self weight. Blue contours represent maximum deflection, on the order of 25.03 cm, which occurs where the convex arches collapse.Figure 15. ANSYS result showing the double arch with lattice structure maintaining its profile. Dotted lines represent the original, undeformed shape.Figure 16.

ANSYS result showing the double arch with lattice structure maintaining its profile. Dotted lines represent the original, undeformed shape.Figure 16. ANSYS result of the double arch with lattice structure showing the deflection contours. Blue contours represent maximum vertical deflection, on the order of 1.1 cm.Figure 17. 3D Concave arch with and without the lattice structure5. ClosureA method has been presented for designing and fabricating deployable structures, with lattice structures as the deployment mechanism. The deployable structures are fabricated from an elastomeric Duraform® FLEX material using selective laser sintering (SLS). The flexible, elastomeric material allows the structures to be folded and compactly stored. For achieving a target surface profile during deployment, lattice structures are utilized both to reinforce the surfaces of flexible structures when they are unfolded and to constrain their expansion when they are pneumatically deployed.

Virtual and physical prototypes provide proof of concept for some preliminary designs with open lattice structures for reinforcement of both convex and concave surface profiles. Future work involves virtual and physical prototyping of deployable structures with closed lattice structures and pneumatic deployment. Prototyping the closed lattice609deployable structures involves several challenges, such as sealing the part with a polyurethane infiltrant to withstand air pressure during deployment and spray coating with a thermoset polymer for stiffness after deployment. Also, an important aspect of the deployment methodology involves prototyping parts that are larger than the SLS build chamber. Currently, large structures are decomposed into smaller parts for fabrication in a nominal build chamber.

Currently, large structures are decomposed into smaller parts for fabrication in a nominal build chamber. Future work involves developing strategies for virtually collapsing (folding, rolling, wadding) the parts in CAD and post-processing the parts to return them to their intended configurations.6. AcknowledgmentsWe gratefully acknowledge financial support from the University of Texas at Austin. Weowe a special thanks to the undergraduate students who participated in this project: Brian Nowotny, Trevor Page, and Brandon Walther.7. References1. 2.3.4.5.6.7.8.9.Gantes, C. J., 2001, Deployable Structures: Analysis and Design, WIT Press, Boston. Pellegrino, S., Ed., 2001, Deployable Structures, International Center for Mechanical Sciences, CISM Courses and Lectures No. 412, Springer-Verlag, New York. Stampfl, J., H. Fouad, S. Seidler, R. Liska, F. Schwager, A. Woesz and P.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDESIGN AND IMPLEMENTATION OF LASER POWDER BED FUSION ADDITIVE MANUFACTURING TESTBED CONTROL SOFTWAREiHo Yeung1*, Keely Hutchinson1,2, Dong Lin21National Institute of Standards and Technology, Gaithersburg, MD 208992Kansas State University, Manhattan, KS 66506AbstractThe National Institute of Standards and Technology developed a facility titled the Additive Manufacturing Metrology Testbed to advance the research in laser powder bed fusion (LPBF) processes. The testbed adopted an open control architecture which allows full access to all key process parameters. Although LPBF control is a very important topic, very little literature can be found on how this is implemented. This paper reviews the testbed control software design and implementation.

Although LPBF control is a very important topic, very little literature can be found on how this is implemented. This paper reviews the testbed control software design and implementation. Scan path planning, galvo motion control, and laser power control are detailed with select highlights. Comparison with commercial machine control software is made, and recent experiments utilizing the advanced features of the testbed control software are also discussed.IntroductionAdditive manufacturing (AM) technology has evolved from its early use in preproduction prototyping to being used regularly to fabricate parts used in many industries, including aerospace, medical, and automotive. It is very useful for creating parts that are too difficult or even impossible to manufacture traditionally.

It is very useful for creating parts that are too difficult or even impossible to manufacture traditionally. One type of AM process is laser powder bed fusion (LPBF) [1], in which a laser is used to selectively fuse a thin layer of metal powder into cross-sections of the intended three-dimensional (3D) structure. This process is repeated after each successive layer of metal powder is spread on top of the completed layer, until the 3D structure is completed. The inherent resulting part quality is determined by many process parameters, such as powder layer thickness, hatch distance, laser spot size, scan path, power, and velocity [2].A major quality issue in additively manufactured parts is pores, that can deteriorate a part’s performance and can cause part failure. Pore formation has been attributed to various phenomena related to the power-speed (P-V) settings, such as keyhole collapse at high laser energy densities [3,4], or insufficient re-melting of adjacent scan vectors [5,6].

The residual stress caused by the very localized rapid heating and cooling cycles in AM is another important issue; it generates distortion and dramatically deteriorates functionality of the additively manufactured parts [7,8]. The thermal history of the AM metal part is also known to affect its microstructure [9,10]. Although AM and specifically LPBF have come a long way and have made significant strides ini Certain commercial equipment, instruments, or materials are identified in this paper in order to specify the experimental procedure adequately. Such identification is not intended to imply recommendation or endorsement by NIST, nor is it intended to imply that the materials or equipment identified are necessarily the best available for the purpose.

* Corresponding author, Email: ho.yeung@nist.gov1455recent years, there are still many opportunities to improve the part quality by optimizing the process parameters and scan strategies.The National Institute of Standards and Technology (NIST) developed the Additive Manufacturing Metrology Testbed (AMMT) to advance monitoring, controls, and metrology research in LPBF [11]. The AMMT control implementation follows the open platform AM process control framework shown in Figure 1. The framework consists of two major parts: the AM software part and the hardware part. The AM software part generates time-stepped digital commands from a computer aided design (CAD) model through slicing, path/power planning, and interpolation steps, and checks the commands created by simulation. Time-stepped digital commands are commonly used in computer numerical control (CNC) where the device status is updated at a regular time interval.

Time-stepped digital commands are commonly used in computer numerical control (CNC) where the device status is updated at a regular time interval. The hardware part executes the digital commands through the synchronized actions of the laser system, powder bed, build chamber, and monitoring devices. Feedback/feedforward control loops can be implemented based on the signals from the monitoring devices.Figure 1: Open platform AM process control framework. The framework consists of four AM software modules and four hardware modules. The interfaces of each software module and functionalities of each hardware module are indicated.In this paper, we focus on the design and implementation of the four AM software modules, namely, slicing, path/power planning, digital interpolation, and simulation (Figure 1). Collectively, we name these four modules as ‘SAM’, or Simple Additive Manufacturing software.

Collectively, we name these four modules as ‘SAM’, or Simple Additive Manufacturing software. SAM provides a reference architecture for an open platform AM control software, which allows researchers to create and test a wide variety of scan strategies, which would otherwise be limited if a commercial machine was used to perform this research. The rest of the paper is divided into three sections: (1) Design and implementation of the SAM; (2) Investigation of scan control on commercial systems; (3) Case studies on the advanced scan strategies implemented with SAM.1456Simple Additive Manufacturing software (SAM)SAM is an open platform AM preparation software developed for use with the NIST AMMT. SAM slices a 3D model into 2D layers, fills the layers with scan paths, creates G-code [12] to describe the scan path, and interprets the G-code path into time-stepped digital commands to drive the testbed. SAM’s interface follows closely the framework from Figure 1.

SAM’s interface follows closely the framework from Figure 1. The SAM interface compared side-by-side with this framework is shown below in Figure 2.Figure 2: SAM interface compared side-by-side with the open platform process control framework.The SAM interface is ‘simple.’ It consists of four clear modules. Each module takes two input files: a parameter file and a data file. The parameter file contains parameters and options required to process the data file. The parameter file is in a simple comma separated variable (CSV) format, all parameters and options are user definable. The STL slicer module takes a STereoLithography (STL) file as the input data file and generates a vertices file which contains vertices of all sliced layers. The G-code generator module takes a vertices file from the previous step as the input data file and generates a G-code file which describes the scan path and power.

The G-code generator module takes a vertices file from the previous step as the input data file and generates a G-code file which describes the scan path and power. The G-code interpreter module takes a G-code file from the previous step as an input data file and generates a digital command file which contains the laser position, power, and diameter command at 10 µs time-steps. This time-stepped digital command file can be sent either directly to the AM controller to build the part, or to the Simulator module to check the file.Each SAM module can also run independently and take a data input file created by a third- party software as long as it follows the same syntax. For example, the SAM G-code interpreter can be used to interpret G-code created by a material extrusion slicer. Obviously, the laser control is missing in the G-code created by a material extrusion slicer but can be easily added, as a G-code file can be created/edited by any text editor.

Obviously, the laser control is missing in the G-code created by a material extrusion slicer but can be easily added, as a G-code file can be created/edited by any text editor. Therefore, the well-defined data file formats, the readiness of third-party software interface, and the easy accessibility of all the control parameters,1457make SAM a truly open platform AM software. Next, we will go through each SAM module in detail.STL SlicerThe first module of the SAM is the STL slicer. STL defines a series of linked triangles to reproduce the surface geometry of the 3D model. STL slicer finds intersect points of the triangles to the layers of constant height parallel to the build plane, and connects the adjacent points on the same layers to create the boundaries for the cross-sections. Figure 3 shows a visual representation of the STL slicing, where Figure 3a-c show a CAD model of a tensile bar, its STL representation, and the parallel slicing planes, respectively.

Figure 3 shows a visual representation of the STL slicing, where Figure 3a-c show a CAD model of a tensile bar, its STL representation, and the parallel slicing planes, respectively. The cross-section boundary created by the intersection of one of the slicing planes with the STL triangles is shown in Figure 3d. Note the boundary consists of only straight-line segments, this is the limitation of the STL representation. If two adjacent line segments are collinear, they can join together to reduce the redundant points as shown in Figure 3e. In order to restore the cylindrical surface, an arc-fitting algorithm is applied as shown in Figure 3f. SAM is able to perform arc motion such as G02 and G03 commands in G-code [12,13]. This greatly improves dimensional accuracy for surfaces with curvature.Figure 3: STL slicing. (a) CAD model of a tensile test bar. (b) STL /triangle representation of the CAD model surface. (c) Parallel slicing planes.

(a) CAD model of a tensile test bar. (b) STL /triangle representation of the CAD model surface. (c) Parallel slicing planes. (d) The boundary of a cross-section outlined by the intersect points (black) resulted from a slicing plane and STL. (e) Collinear segments joined. (f) Arc-fitting to restore the surface with curvature.Parameters such as the layer heights, starting and ending layers, and arc-fitting tolerance can be customized in the parameter file for this module. The output of this module is a binary file containing vertices of all sliced layers, with a mark-up to specify whether the two adjacent points are connected by a straight line or an arc. If it is an arc, the center and radius of the arc are also specified.

If it is an arc, the center and radius of the arc are also specified. This vertices file provides all necessary geometric information for the next step, the G- code generation.1458G-Code GeneratorThe next module of SAM (Figure 2) is the G-code generator, which generates AM G-code that describes the scan path and power used for each layer. The AM G-code is a slightly modified version of the RS-274 standard [14]. It adds keywords ‘L’ for laser power, and ‘D’ for laser diameter control. An example is shown in Figure 4, where the laser is commanded to go linearly (G01) to coordinate (X=5, Y=0) with speed (F) = 1000 mm/s, laser power (L) = 200 W, and laser diameter (D) = 2 times of the default setting. The L and D keywords allow the nominal laser power and diameter to be set for each scan line, which allows more flexibility in scan strategy implementation. In addition to linear motion (G01), arc motions (G02 and G03) and dwell (G04) are also implemented following the RS-274 standard.

In addition to linear motion (G01), arc motions (G02 and G03) and dwell (G04) are also implemented following the RS-274 standard. A new capability is also developed to realize spiral paths in Figure 5a.Figure 4: An example line of AM G-code defining a move of the laser.There are many options for the path and power planning in SAM; they are summarized in Figure 5. Four sample hatching patterns are shown in Figure 5a. In a raster pattern, evenly spaced parallel lines are used to fill the region being built. In a concentric pattern, the region is filled by a series of concentric contours. The concentric spiral pattern is similar, but it spirals in with a continuous scan path. The shift spiral pattern is a moving circular path, which leaves a fully melted stripe behind. Different hatching patterns can be applied between layers. There are more detailed parameters that can be edited to more fully describe how these paths are carried out.

Different hatching patterns can be applied between layers. There are more detailed parameters that can be edited to more fully describe how these paths are carried out. Certain patterns may not fit completely to the region being built, so a ‘power mask’ is implemented to shut off the laser power instead of disrupting the scan motion when it goes beyond the part boundary. Therefore, these hatch patterns can be applied to any shape.Two examples of hatching power modes are shown in Figure 5b. On the left is the interleaved power mode, which scans every other line on the first pass, and the remaining lines on the second pass. On the right is alternative power mode, in which the laser power is alternated between two power levels on sequential scan lines. These modes can be controlled in detail using the parameters file.

On the right is alternative power mode, in which the laser power is alternated between two power levels on sequential scan lines. These modes can be controlled in detail using the parameters file. There are also options for the island strategy that allow the region to be divided into rectangles of any size, and scanned in different sequences with different inter-island rotation angles, as shown in Figure 5c. The same layer can be scanned multiple times, with different power levels, laser diameters, and hatch space settings, as shown in Figure 5d. There are also options in SAM to set layer-wise scan strategies, such as different nominal power and speed, and interlayer rotation angle for different layers. Many of these options can also be combined to create highly customized scan strategies for different research needs. The output of this module is an AM G- code file. This output can also be further customized by the user before its interpretation.1459Figure 5.

The output of this module is an AM G- code file. This output can also be further customized by the user before its interpretation.1459Figure 5. Examples of path/power planning modes available for use in SAM. (a) Hatching patterns. (b) Hatching powers. (c) Island sequence. (d) dual scans.G-Code InterpreterAM G-code contains information of destination, nominal speed, nominal power, and nominal spot size. It needs to be converted into a time-stepped digital command for the AM controller to execute. This conversion is carried out in the G-code interpreter module. The time- stepped digital command file in SAM is defined as an n x m numerical array as shown in Figure 6a, where n is the number of time steps in 10 μs increments, and m is the number of control parameters. The command format is based on the xy2-100 protocol [18] for galvo position (X, Y), but extended to laser power (L), laser diameter (D), and trigger for monitoring devices (T).

The command format is based on the xy2-100 protocol [18] for galvo position (X, Y), but extended to laser power (L), laser diameter (D), and trigger for monitoring devices (T). One line of the commands is executed by the AM controller at every 10 μs (Figure 6b), which sends the update to each controlled device (Figure 6c).Figure 6. Time-stepped digital command and its execution. (a) Command format. Each column is a control parameter. (b) Command execution. Every 10 µs a line is sent for execution. (c) Devices controlled.1460To interpret G-code into time-stepped digital commands, three interpretation modes are defined. (1) Velocity mode for motion control. (2) Path mode to handle the beginning and end of a scan. (3) Power mode to synchronize the laser power to motion. Figure 7 shows the three velocity modes implemented in SAM. Step velocity (Figure 7a) is the easiest to program but it requires an infinite acceleration, which no physical system can deliver.

Figure 7 shows the three velocity modes implemented in SAM. Step velocity (Figure 7a) is the easiest to program but it requires an infinite acceleration, which no physical system can deliver. Trapezoidal velocity (Figure 7b) allows time for the velocity to pick up, but it assumes the required acceleration can be reached instantaneously. For the galvo (a DC motor), the torque/acceleration is proportional to the current flow through the armature. The DC motor is a second-order system, it does not allow a jump in current. Therefore, the time derivative of the acceleration, which is known as jerk, cannot be infinite. An infinite acceleration or jerk will cause deviation of the actual velocity from the commanded velocity and hence cause the following error.

An infinite acceleration or jerk will cause deviation of the actual velocity from the commanded velocity and hence cause the following error. Jerk-limited velocity (Figure 7c) gives the best spatial and temporal accuracy [15], but it is also the most computational intensive to program.Figure 7: Position, velocity, and acceleration graphs for (a) step velocity, (b) trapezoidal velocity, and (c) jerk-limited velocity profile. The jerk-limited velocity here is implemented with a sinusoidal acceleration.Figure 8 shows the three path modes defined in SAM. The exact stop mode (Figure 8a) stops motion exactly at the end of each scan with maximum allowable deceleration, and the laser power is switched off when the motion is stopped. If there is a subsequent scan, the motion will start again immediately with maximum allowable acceleration until it reaches the nominal speed (set in AM G-code by keyword F), or until it needs to decelerate again.

The constant build speed mode (Figure 8b) keeps the programmed speed at the end of the current scan, and adds a colinear deceleration path to stop the galvo. It also adds a colinear acceleration path to the start point of the next scan, to speed up the galvo to the programmed speed. The laser power is turned off during these deceleration/acceleration paths. It is named constant build speed mode since whenever the laser power is on, the scan speed is constant. The continuous mode (Figure 8c) matches the ending velocity of the scan with the beginning velocity of the subsequent scan. To do so, an arc is used to join two adjacent scans (hatch lines). The laser power is kept on, hence the meltpool is also continuous [16].1461Figure 8: Speed plot for the three path modes. The thicker lines indicate where the power is on, the pseudo color indicates the speed. (a) Exact stop. (b) Constant build speed.

The thicker lines indicate where the power is on, the pseudo color indicates the speed. (a) Exact stop. (b) Constant build speed. (c) Continuous.Three power modes are also defined in SAM: constant power, constant power density, and thermal adjusted power. Constant power mode keeps the power constant at the nominal level (set in AM G-code by keyword L) for the whole scan line. Constant power density mode keeps the power/speed ratio (power density) constant within a scan line. The thermal adjusted power mode compensates for the local variation of thermal property by changing the laser power. It also provides a means for model-based feedforward control of the laser power. Users can also develop their models as plug-ins.The AM G-code interpretation and time-stepped digital command file enables continuous laser power control, complicated scan path implementation, and model-based scan strategy optimization.

The use of the time-stepped digital command also makes the AM process fully transparent. The digital command file alone contains all the information required to build a part, and hence supports the cross-platform operation. It will be shown later that the digital commands intercepted on a commercial system can also be executed on AMMT.SimulatorAM is a very expensive process. Parts usually take hours if not days to build. It is crucial to examine the build files to make sure there are no design errors. The simulator reads the time- stepped digital commands, simulates the scan, and displays the result graphically for the user to examine. Figure 5, Figure 7, and Figure 8 are prepared with the Simulator. More sophisticated simulations, such as the galvo dynamic response and meltpool behavior, are under development.Commercial SystemsFor an LPBF AM system, the heating laser is guided to the build platform by a pair of mirrors driven by galvo motors.

Figure 9 shows a simplified schematic of the scan control. The AM scan paths, which are a series of x-y position pairs, are created by AM preparation software. These positions are then interpolated into time-stepped digital commands and transmitted through digital communication lines following the xy2-100 protocol. The digital/analog (D/A) receiver then converts the digital commands received to analog voltages to move the mirrors, and thereby moving the laser beam through the galvo driver board. On most commercial machines, however, the system described in Figure 9 acts as a black box. The user has limited control over how the1462scan path is created and interpolated. To understand the scan strategy and control on these commercial systems, one approach is to tap into the xy2-100 digital communication line (Figure 9) and intercept the digital command. From these digital commands, the scan strategy can be reconstructed.Figure 9: The schematics show AM scan control.

From these digital commands, the scan strategy can be reconstructed.Figure 9: The schematics show AM scan control. The red arrow shows where the digital command signal can be intercepted.Figure 10 shows the scan strategy read from a commercial LPBF AM machine, where single tracks are scanned, and the digital commands are intercepted. The programmed scan path is shown in Figure 10a, and the scan path constructed from the captured digital signal is shown in Figure 10b. The x-displacement, x-y speed, and acceleration are plotted in Figure 10c. The nominal scan speed is 200 mm/s, but it can be seen there are jumps with a speed of approximately 5000 mm/s and acceleration of 5x108 mm/s2. The acceleration requires increasing the speed from 200 mm/s to 5000 mm/s in one 10 µs time step; therefore a step velocity profile was applied for the motion control. No galvo system can deliver this acceleration; therefore the following error must occur at the jump.

No galvo system can deliver this acceleration; therefore the following error must occur at the jump. Nevertheless, it should be pointed out that the following error does not necessarily cause a build issue, as long as the laser power is not on.Figure 10: Scan strategy read from a commercial LPBF AM machine for single tracks. (a) Programmed scan path. (b) Scan path constructed from digital command intercepted. (c) X-displacement, x-y speed, and x-y acceleration.Figure 11 shows the scan strategy read from a commercial LPBF AM machine of a different model, where a square pad was scanned (Figure 11a). The initial infill and the contour regions are shown in Figures 3b and 3c, respectively, alongside the speed and acceleration plots. The laser power is recorded simultaneously, and the path taken with the laser on is plotted in red. The infill is programmed at 750 mm/s and the contour at 400 mm/s. Again, a step velocity profile1463is applied for the motion control.

The infill is programmed at 750 mm/s and the contour at 400 mm/s. Again, a step velocity profile1463is applied for the motion control. For the infill, jumps are observed at the beginning/end of every hatch line (Figure 3b), but they do not happen immediately before/after the laser power is turned on/off. An extension (overshoot) of the path is inserted at the beginning/end of the scan to allow the system to recover from the following error caused by the jump. Therefore, during the build (when the laser power is on) the scan speed remains constant. This strategy is known as skywrite on some commercial machines. This is very similar to the constant build path mode of SAM, except this extension path is at a constant speed, not an accelerate/decelerate path that SAM uses.However, the contour is scanned at a constant speed without overshoot.

It is uncertain if the sharp turn at the corner can be followed, but the effect on part geometry is likely minimal since the area will be melted again by the infills. Step velocity simplifies the interpretation, but a significant following error is unavoidable. The potential build issues can be avoided by carefully designed/calibrated scan strategies such as the skywrite here, but this also imposes restrictions on the scan strategy design.Figure 11: Scan strategy read from a commercial LPBF machine for a square pad. (a) Scan path constructed from digital command intercepted. The path with laser power-on is plotted in red. (b) Enlarged infill path for the shaded region in (a). The end-of-scan jumps are marked by arrows on the x-y speed plot on the left. (c) The contour, no overshoot is observed.The galvo position read by tapping into the xy2-100 digital communication line is the commanded position.

(c) The contour, no overshoot is observed.The galvo position read by tapping into the xy2-100 digital communication line is the commanded position. The actual galvo position can be measured by reading the galvo position encoder feedback (Figure 9). However, the galvo encoder signal is not easily accessible on commercial machines. Since AMMT has already built in the capability to capture the galvo encoder position, a commercial AM controller can be used to drive AMMT. The behavior of this commercial AM controller is studied by comparing the commanded and measured galvo positions (measured by the galvo position encoder).1464Figure 12a shows how an 8 mm square is scanned on AMMT driven by a commercial AM controller. The controller software provides a graphic user interface (GUI) to define the scan path, speed, and power. Both the commanded and measured positions were captured and plotted in Figure 12a in blue and orange, respectively.

Both the commanded and measured positions were captured and plotted in Figure 12a in blue and orange, respectively. The position, speed, and acceleration plots in the figure show that a step velocity profile is used for the command, and the sharp corner of the commanded position could not be followed. This is shown in the speed plot in Figure 12a. It is surprising to see that during the turn, the scan is commanded to idle for a short moment after the current line stops and before the next line starts, while the power is still on. It is marked by ‘X stop’ and ‘Y start’ in the position plot in Figure 12a. This idle is believed to be inserted by the controller to ensure that the current line is completed before the next line starts, to maintain a sharp corner (geometric accuracy), although this results in an overheated region at the corner. Figure 12b shows the same square scanned with SAM software and AMMT controller. The ramp velocity profile is used for the motion control.

Figure 12b shows the same square scanned with SAM software and AMMT controller. The ramp velocity profile is used for the motion control. The position and velocity profiles are better followed. In this case, the improved temporal accuracy allows us to start the Y scan immediately after X stops. That reduces build time and prevents overheating.Figure 12: Commanded and measured scan position, velocity, and acceleration for (a) commercial AM controller, and (b) SAM controller. An 8 mm square was scanned. The commanded positions are plot in blue, the measured in orange. Xc, and Xm are the commanded and measured positions, respectively.Although it is a very important topic and directly impacts the build quality, there is very little literature that can be found on how the LPBF AM control is implemented. We could only conduct very limited studies based on the equipment and tools available. Nevertheless, our goal1465here is to promote open platform AM process control.

We could only conduct very limited studies based on the equipment and tools available. Nevertheless, our goal1465here is to promote open platform AM process control. We already presented our control framework and the SAM software developed following this framework. In the next section, we will use a few case studies to demonstrate its benefits.Advanced Scan Strategy Case StudiesThe AM control framework (Figure 1) and the open platform modular design (Figure 2) allow a continuous improvement of SAM software to add functions and features per research needs. In this section, we present case studies on the advanced scan strategies developed with SAM. The development of these scan strategies, however, is inspired by many excellent AM research works.Spiral Scan StrategySpiral scan strategy is developed to research the effect of an unconventional, non-linear scan strategy on meltpool area and the resulting microstructure.

The design of the continuous circular spiral scan path can be seen in Figure 13a. It starkly contrasts the conventional raster scan path in Figure 13b. This unconventional scan strategy is made possible through the G22 implementation in SAM. 17-4 stainless steel rectangular parts of 10 mm in length, 5 mm in width, and 10 mm in height were built with both raster and spiral scan strategies. The parts built with the spiral scan strategy show a maximum of 54% increase in hardness and a distinct microstructure. It is speculated the much larger meltpool observed and the reheating effect by the spiral scan strategy significantly changed the solidification process. The spiral scan provides a possible means to control the microstructure of the additive manufactured material. More detail of this work can be found in [17].Figure 13: (a) Spiral scan strategy. (b) raster scan strategy. The arrows indicate the scan directions.

More detail of this work can be found in [17].Figure 13: (a) Spiral scan strategy. (b) raster scan strategy. The arrows indicate the scan directions. The hatch space (space between the circles or the lines) is 100 µm.Three-Dimensional Scan StrategiesTwelve nickel alloy 625 rectangular parts of 10 mm in length, 10 mm in width, and 5 mm in height were built. The goal of this work was to compare the effect of scan strategy on part quality. A different scan strategy is used for each part (Table 1). These scan strategies were composed with the SAM using the combination of its standard options in the G-code generator and G-code interpreter modules. For example, part 8 used a combination of island strategy, exact stop path mode, and constant density power mode. The process was monitored in situ by a high- speed camera coaxially aligned with the heating laser beam (Figure 6c). The galvo encoder position1466was captured.

The process was monitored in situ by a high- speed camera coaxially aligned with the heating laser beam (Figure 6c). The galvo encoder position1466was captured. All the parts were successfully built, and the time-stepped digital command and monitoring data were published [18] and used by many scan strategy studies. The parts were also measured with X-ray Computed Tomography (XCT) for porosity, and correlations between the pores and the meltpool images were attempted.Table 1. The twelve three-dimensional scan strategies studiedFigure 14 compares the meltpool from parts 7 and 8, that used constant power and constant power density scan strategies, respectively. For each scan strategy, four sample meltpool images are shown on the right, while the speed, power, and meltpool image area are plotted on the left. The meltpool image area is measured by thresholding the 8-bit grayscale meltpool image at intensity level of 100.

The meltpool image area is measured by thresholding the 8-bit grayscale meltpool image at intensity level of 100. The constant power density was implemented with Equation 1, where Lo is the nominal laser power in W, L is the applied laser power in W, V is the instantaneous speed in mm/s, Vo is the nominal speed in mm/s, and C is a unitless weighting factor between 0 and 1. C = 0.5 is used to maintain a minimum laser power level by varying only a portion of the power to the speed since power should not drop below a threshold where the powder will not melt at any speed. It can be clearly seen from Figure 14 that the constant power density scan strategy kept the meltpool more constant when the speed slows down.L = (V/Vn) · C · Lo + (1-C) · LFigure 14: Meltpool area plotted together with laser speed and power shown on the left, and meltpool images at locations marked by an ‘x’ shown on the right. (a) Constant laser power.

(a) Constant laser power. (b) Constant power density1467(1)The constant power density scan strategy requires continuous laser power variation (intra- vector laser power control), which is made possible by the SAM G-code interpreter. The XCT results show the porosity decreased from 0.39% for the part built by constant power to 0.11% for the part built by constant power density strategy. It is believed the constant power density strategy reduces the concentration of laser energy at the end of the scan and hence prevents the formation of keyhole pores.Geometric Conductivity FactorWhen a part is built using LPBF, there are areas that have powder, melted metal, and solid metal. These areas have different thermal conductivities depending on the state of the metal and the part geometry. This can be explained by Figure 15, where voxels 1 – 4 have different conductivity, depending on whether their neighborhood is ‘more’ solid or powder when it is being built.

This can be explained by Figure 15, where voxels 1 – 4 have different conductivity, depending on whether their neighborhood is ‘more’ solid or powder when it is being built. A Geometric Conductivity Factor (GCF) is developed to quantify this effect. The GCF model is created solely based on the time-stepped digital command. The computer algorithm follows the command to reconstruct a 3D structure by ‘melt’ voxels as in Figure 15, and assigns each melt voxel a GCF value as the voxel is created (melted).Figure 15: 3D structure reconstructed through digital command. The darker voxels are where the laser power is programmed to be on.Using a Geometric Conductivity Factor (GCF) scan strategy, the different conductivities due to the geometric shape and build sequence can be considered. The laser power can be altered to account for these differences. Figure 16 shows an example, where a GCF model is created from the digital command of overhanging structure.

The laser power can be altered to account for these differences. Figure 16 shows an example, where a GCF model is created from the digital command of overhanging structure. The GCF values near the boundaries and the overhang regions are smaller since these regions are surrounded by powder (Figure 16a). The powder has a much lower conductivity than solidified metal [19]. The GCF is then used to adjust the laser power, as shown in Figure 16b. This is a typical example of model-based thermal adjusted power mode in SAM. The part built with the GCF scan strategy demonstrated a significant reduction in surface roughness in the overhang regions. More details on the GCF model construction and its effect on part qualities can be found in [20].1468Figure 16: GCF scan strategy construction. (a) GCF model. (b) GCF optimized scan strategy for one of the layers.

(a) GCF model. (b) GCF optimized scan strategy for one of the layers. Laser power is represented by the pseudo color.Residual Heat FactorThe residual heat factor (RHF) is created to account for the dynamic reheating effect. This is an effect that occurs because previous scan lines will have residual heat, affecting the thermal condition of the current scan position. The dynamic reheating effect often results in inconsistent meltpool morphology. To counteract this, a value called RHF is assigned each scan position (points in the time-stepped digital command) to ‘quantify’ the residual heat effect, and the laser power is adjusted based on the RHF value to reduce local variability of meltpool area. This is another example of model-based thermal adjusted power mode. The RHF concept is based on the proximity analysis feature of SAM [21], which determines the neighborhood scan points in a predefined temporal and spatial window as shown in Figure 17.

The RHF concept is based on the proximity analysis feature of SAM [21], which determines the neighborhood scan points in a predefined temporal and spatial window as shown in Figure 17. The temporal window defined by the elapsed time T and spatial window defined by a radius R are drawn in the figure on the scan path. Each dot on the path is a position in the time stepped digital command. The RHF for the scan point i can be formulated as Equation 2, where dki = distance of point i from point k, tki = time elapsed between point i and point k, Lk = laser power level at point k, vk is the velocity at point k, and S is the set of points scanned within time T and within distance R from point i.RHFi = ∑ ( 𝒌∈𝑺𝑹−𝒅𝒌𝒊 𝑹)2(𝑻−𝒕𝒌𝒊 𝑻)𝑳𝒌 𝒗𝒌(2)Figure 17: RHF model construction. Each dot on the scan path is a point in the time-stepped digital command.

Each dot on the scan path is a point in the time-stepped digital command. Elapsed time < T is indicated by a purple dotted line and distance < R indicated by a red circle, where T = 40 time-steps and R = 250 µm in this example.1469RHF in Equation 2 assumed the residual heat effect from a previous point k to the current scan point is inversely proportional to the elapsed time and square of the distance. A different heat dissipation model can be used. The RHF is then normalized and used to scale the laser power. The RHF is optimized by the value of T and R, which can be done by simulation, machine learning, or experimentally. An experimental approach to RHF optimization can be found in [22]. Figure 18 shows an example, where a 3 mm x 2 mm rectangular pad was scanned without and with the RHF optimization, and the meltpool area maps are plotted in Figure 18a and Figure 18b, respectively.

Figure 18 shows an example, where a 3 mm x 2 mm rectangular pad was scanned without and with the RHF optimization, and the meltpool area maps are plotted in Figure 18a and Figure 18b, respectively. The variability of meltpool area is reduced by 27% (measured by the standard deviation of the meltpool area) by the RHF optimization, and this reduction can be clearly seen by comparing Figure 18a and Figure 18b.Figure 18: Meltpool area map for a rectangular pad scan. The meltpool area is represented by pseudo color. (a) Without RHF optimization. (b) With RHF optimization.Machine LearningThe RHF concept can also be used to formulate a data-driven meltpool area predictive function, such as Equation 3, where 𝑆𝑖 is the meltpool size for scan point i in the time-stepped Δt represent the digital command, and Li is the laser power, vi is the scan speed, 𝜃𝑖 cumulative residual heat effect based on spatial difference ∆d and time difference ∆t.

They can be as in the RHF definition represented as 𝜃𝑖 (Equation 2). A different representation can also be used, such as in our meltpool prediction based ∆𝑡 is put as scan strategy study [23], in which 𝜃𝑖 ∑ (𝑒𝑎𝑡𝑡𝑘𝑖)𝛥𝑑 and 𝜃𝑖𝑹−𝒅𝒌𝒊 𝑹𝑳𝒌 𝒗𝒌𝑻−𝒕𝒌𝒊 𝑻𝑳𝒌 𝒗𝒌∆𝑑 = ∑ ( 𝒌∈𝑺∆𝑡 = ∑ ( 𝒌∈𝑺)2and 𝜃𝑖)∆𝑑 is dropped to prevent overfitting, and 𝜃𝑖𝑳𝒌 𝒗𝒌, where 𝑎𝑡 is a constant which can be determined from the experiment.𝒌∈𝑺𝑆𝑖 = 𝑓(𝐿𝑖, 𝑣𝑖, 𝜃𝑖∆𝑡, 𝜃𝑖∆𝑑) (3)Once a meltpool area prediction model is trained, it can be used to adjust the laser power in the time-stepped digital command until the calculated meltpool area is within the objective, or the power constraint reached. A combination of objective meltpool area and constraint of power was tested in [23]. The results are shown in Table 2 together with the original scan strategy. The original scan strategy used in this exercise is the island concentric, continuous, and constant power density strategy from #6 in Table 1.

These are excellent works, but there is a lack of experimental verification. Most commercial systems do not accept user-optimized scan strategies. This is also true for testbeds developed with commercial galvo controllers; user customization can only stay at the scan line level since there is no access to in-line control. The NIST AM framework (Figure 2) uses time-stepped digital commands as process control files, which supports user customization down to each scan point. The pointwise control in the time- stepped digital command (Figure 4) by its nature supports the machine learning-based optimization method very well.Conclusion and Future WorkAn additive manufacturing (AM) control software, Simple AM (SAM), is developed. SAM consists of four modules: slicing, planning, interpolation, and simulation. The slicing module slices the STL file and outputs a vertices file that contains the vertices of all the layers sliced.

SAM consists of four modules: slicing, planning, interpolation, and simulation. The slicing module slices the STL file and outputs a vertices file that contains the vertices of all the layers sliced. An AM G- code standard is proposed, adding the laser power and laser spot diameter control on top of the position and speed control of conventional G-code. The planning module generated scan path and power to cover the regions described by the vertices file and output the results as AM G-code. A time-stepped digital command file format is also defined, consisting of galvo X-Y position, laser power, laser diameter, and periphery device trigger updates in 10 µs incremental time steps. An AM G-code interpreter is developed to convert AM G-code to time-stepped digital command based on different velocity profiles, path modes, and power modes. This is also known as interpolation.

An AM G-code interpreter is developed to convert AM G-code to time-stepped digital command based on different velocity profiles, path modes, and power modes. This is also known as interpolation. The time-stepped digital command file can be sent directly to execute on the AM machine, or verified on the simulation module.Laser powder bed fusion additive manufacturing is a very delicate process. A small deviation from the optimum condition could trigger defects and deteriorate the part quality. Such deviation can be caused by inadequate process control, such as a power concentration at the end of the scan, or a change in the thermal properties, such as a change of conductivity over an overhang region. We demonstrated how these situations are handled through different case studies. One common requirement in these studies is the in-line power control, which is made possible by the SAM interpreter and the time-stepped digital command.

One common requirement in these studies is the in-line power control, which is made possible by the SAM interpreter and the time-stepped digital command. The time-stepped digital command allows a point-wise control of the AM process, versus the line-wise control which is commonly1471used on AM machines today. Moreover, the time-stepped digital command also facilitates model- based scan strategy optimization.One uniqueness of our work is the jerk-limited motion control; it improves both spatial and temporal position accuracies, hence enabling an accurate laser power-position synchronization. This established the foundation of advanced AM control. The motion control on commercial systems was compared. It is found that the step velocity profile was used for the motion control on all three systems we evaluated. It is interesting to see how the overshoot path and idling time are used to prevent the following errors and improve spatial accuracy.

It is interesting to see how the overshoot path and idling time are used to prevent the following errors and improve spatial accuracy. These limit the flexibilities in scan strategy but allows a very computationally efficient path interpolation. Nevertheless, computational efficiency may no longer be an issue on a modern numerical controller with improved computational performance. The same is true for the data storage; the large file size of the time-stepped digital command is less and less of a concern.There are many opportunities to improve the field of laser powder bed fusion (LPBF) additive manufacturing. More research into the laser scan strategies has the potential to develop repeatable ways to produce high-quality parts with minimal defects, controlled residual stress or microstructure, and improved speed and efficiency of material consolidation. The open platform software gives researchers the maximum flexibility possible and maximum control over process parameters.

The open platform software gives researchers the maximum flexibility possible and maximum control over process parameters. Many built-in features of SAM have not been tested/optimized. The machine learning scan strategy optimization also seems very promising. These are the directions we are continuously working toward.Reference[1] Badiru AB, Valencia VV, Liu D, editors. Additive Manufacturing Handbook: Product Developmentfor the Defense Industry. 1st ed. CRC Press; 2017. https://doi.org/10.1201/9781315119106. [2] Oliveira JP, LaLonde AD, Ma J. Processing parameters in laser powder bed fusion metal additivemanufacturing. Materials & Design 2020;193:108762. https://doi.org/10.1016/j.matdes.2020.108762. [3] King WE, Barth HD, Castillo VM, Gallegos GF, Gibbs JW, Hahn DE, et al. Observation of keyhole-mode laser melting in laser powder-bed fusion additive manufacturing. Journal of Materials Processing Technology 2014;214:2915–25.[4] Gong H, Rafi K, Gu H, Starr T, Stucker B.

sign And Integration Of A Laser-Based Material Deposition SystemJ. Laeng@Jamaluddin Abdullah*, Frank Liou**, M.N. Mohamad Ibrahim♣ and W.M. Wan Muhammad*School of Mechanical Engineering, Universiti Sains Malaysia - Engineering Campus, 14300 Nibong Tebal, Seberang Perai, Penang, MALAYSIA ♣School of Chemical Sciences, Universiti Sains Malaysia, 11800 Penang, MALAYSIA **Department of Mechanical Engineering, University of Missouri-Rolla MO 65409-1350, USAABSTRACTThis paper aims to demonstrate the design process of an integrated five-axis Laser-Based Material Deposition (LBMD) system for rapid prototyping application. Several design evaluation methods are selected and applied to the design of the system. A three-dimensional graphical simulation software package was used as a decision making aid and as an analysis tool in the design process.

A three-dimensional graphical simulation software package was used as a decision making aid and as an analysis tool in the design process. Hardware integration of a five-axis computer numerical controlled (CNC) vertical milling machine, a 2.5 KW Nd:YAG laser and a linear table is discussed. A brief introduction to the system software and control architecture is also summarized. Some important design issues and considerations specific to Laser Based Material Deposition process are suggested.Keywords: Design methodology, System Design and Integration, System Simulation, Laser Based Material Deposition, Collision Detection, Tool Path Verification1. INTRODUCTIONSince the introduction of the first commercial stereolithography apparatus in 1986, many rapid prototyping (RP) systems have been developed for the ever growing industrial demands for shorter product development cycle time and at significantly reduced costs [1].

Most of the commercial RP units at that time however, were either limited to non-metal fabrication with minimal accuracy and precision or constraint to produce relatively simple parts for physical visualization of design concepts. Some of the developing systems were capable of producing direct metal prototypes. However, the commercial units still require polymer RP molds from which metal castings are formed [2].The increasing complexity of certain parts geometry imposed by many advanced applications making the production of highly accurate functional model directly from CAD data very difficult or impossible using the available RP systems. Arising from this, many systems have been proposed. The LENS (Laser-Engineered Net Shaping) process has demonstrated the feasibility to produce near net shape metal part [3,4].

Arising from this, many systems have been proposed. The LENS (Laser-Engineered Net Shaping) process has demonstrated the feasibility to produce near net shape metal part [3,4]. Others have tried to utilize the existing laser cladding technique, an established laser surface coating of metallic parts, to produce metal parts by layer additive approach [2,5].11This paper describes the design and integration process at the early development stage of a Laser Based Material Deposition process at the University of Missouri-Rolla, USA. A few design evaluation methods, including a graphical simulation software were employed to assist in decision making. The LBMD system consists of a 5-axis CNC machine, Nd:YAG laser and a linear table put together to work as an integrated system.2 DESIGN AND INTEGRATION PROCESSA survey on Laser Based Material Deposition was conducted in order to understand the process and other related issues [6].

Three design concepts were proposed as shown in figure 1, figure 2 and figure 3. The first design as shown in figure 1, integrates a 3-axis CNC milling (X,Y and Z), a 2-axis rotary table (tilting about X, rotating about Z) and a single axis linear actuator (slide up and down along Z axis) that carries the laser and powder delivery nozzle. A rigid support structure is constructed above and spans across the CNC assembly. The laser powder nozzle is attached to this support bar. This design concept provides a stable and rigid support for laser and powder assembly and thus minimizes vibration.Laser NozzleLaser NozzleCNC Machine SpindleRotary TableCNC Machine SpindleRotary TableFigure 1 : Design Concept I.Figure 2: Design Concept II.In design concept II as shown in Figure 2, a different design concept is proposed. The laser and powder delivery nozzle is attached to the CNC machine head assembly besides the machine spindle.

The laser and powder delivery nozzle is attached to the CNC machine head assembly besides the machine spindle. This design integrates a 3-axis CNC milling machine, 2-axis rotary table and a single axis laser and powder nozzle. No support structure as in design I need to be constructed. Initially, this design is thought to be better because it eliminates the support structure for the laser and powder nozzle and does not require additional motion axis for the system. However, the cost of putting the laser nozzle to the CNC machine head assembly will be significantly large. Vibration from the machine spindle during machining process could cause misalignment of laser optic just next to it. Laser optics are also exposed to contamination from metal chips and also coolant.

Vibration from the machine spindle during machining process could cause misalignment of laser optic just next to it. Laser optics are also exposed to contamination from metal chips and also coolant. Chances of collision between work piece and the machine spindle and also between work piece and laser nozzle increases during rotation or tilting of work piece.In concept Design III as shown in Figure 3, a cantilevered support with retraction and12extension arm, where laser and powder nozzle is attached to, is proposed for the system. The laser and powder nozzle can be extended to the work area during metal deposition and they can be retracted to a safe location during machining process. This feature is hoped to protect the laser optics from metal chips and contaminants within the vicinity of the work area. However, accuracy could be jeopardized due to deflection in the cantilever arm and cumulative displacement error from the frequent retraction and extension of the cantilever arm.

However, accuracy could be jeopardized due to deflection in the cantilever arm and cumulative displacement error from the frequent retraction and extension of the cantilever arm. Also, to have this feature, two linear actuators are needed and considering the load and moment the beam will carry, achieving a higher displacement accuracy (0.001”) becomes impossible. With the design configuration in concept design III, the total axis of control would be increased to 7. This is not desirable because it makes integration and controls of the system more complex. Synchronization and motion control of all the individual components is critical to part quality. Therefore, a simple system, but capable of good motion control is desirable.Retracting ArmLaser NozzleCNC Machine SpindleRotary TableFigure 3: Design Concept IIITwo methods were used in design evaluation process. The first method utilizes concept evaluation matrix to compare each design to a selected datum concept.

The first method utilizes concept evaluation matrix to compare each design to a selected datum concept. This method provides three levels of comparison: “better than (+) ”, “the same as (S) ” or “less/worst than ( - )” the datum. A set of evaluation criteria was established. Table 1, Table 2 and Table 3 shows the evaluation matrix for each concept. Based on this evaluation, design concept I is the most feasible option. The second method uses a scale from 1 to 5 to further evaluate the design. Table 4 shows the scores for each design. This method assigns quantitative score to each concept and we can rank the concept, as well as each criterion from the highest to the lowest score. The highest score means the best and the lowest is the least/worst. From the second evaluation method, concept I carries the highest score. Concept design I is finally selected for further analysis.

The LBMD system would fabricate parts by combination of material addition (laser material deposition) and material removal (CNC milling to correct deposition errors).13Table 1: Concept Evaluation Matrix Using Conceptual Design I as DATUMConceptCriteria Positioning accuracy of the system Repeatability of the system Ease of system operation Ease of maintenance of laser and powder Protection of laser alignment from vibration Protection of nozzle and laser from Simplicity of Integration Simplicity of Synchronization User safety from Laser Ergonomics (user movement and interaction) Minimum axis of control Working volume Rigidity and stability of nozzle support Overall estimated cost for the system Adaptability to multi material deposition Online process control implementation Score1 D A T U MD A T U MD A T U M2 S S + _ _ _ + + S S + S + _ _ + 6+, 5-, 5STable 2: Concept Evaluation Matrix Using Conceptual Design II as DATUMConceptCriteria Positioning accuracy of the system Repeatability of the system Ease of system operation Ease of maintenance of laser and powder Protection of laser alignment from vibration Protection of nozzle and Simplicity of Integration Simplicity of Synchronization User safety from Laser Ergonomics (user movement and interaction) Minimum axis of control Working volume Rigidity and stability of nozzle support Overall estimated cost for the system Adaptability to multi material deposition Online process control implementation Scorelaser from1 S S _ + + + _ _ S S _ S _ + + S 5+, 6S, 5-2 D A T U MD A T U MD A T U M143 _ _ S S + _ S S S S S _ _ S _ _ 1+, 7-, 8S3 _ _ _ + + + _ _ S S _ S _ S + _ 4+, 4S, 8-Table 3: Concept Evaluation Matrix Using Conceptual Design III as DATUMConceptCriteria Positioning accuracy of the system Repeatability of the system Ease of system operation Ease of maintenance of laser and powder Protection of laser alignment from vibration Protection of nozzle and laser from contaminants Simplicity of Integration Simplicity of Synchronization User safety from Laser Ergonomics (user movement and interaction) Minimum axis of control Working volume Rigidity and stability of nozzle support structure Overall estimated cost for the system Adaptability to multi material deposition Online process control implementation Score1 + + + S S S S S S S + S S + + S 6+, 10S2 + + + _ _ _ + + S S + S + _ _ + 8+, 5-, 3STable 4: Concept Evaluation Using Quantitative Scale (scale 1 – 5)ConceptCriteria Positioning accuracy of the system Repeatability of the system Ease of system operation Ease of maintenance of laser and powder Protection of laser alignment from vibration Protection of nozzle and laser from contaminants Simplicity of Integration Simplicity of Synchronization User safety from Laser Ergonomics (user movement and interaction) Minimum axis of control Working volume Rigidity and stability of nozzle support structure Overall estimated cost for the system (least is best) Adaptability to multi material deposition Online process control implementation Score1 3 3 4 4 5 5 4 4 3 3 4 4 4 5 4 4 582 5 4 3 2 1 2 2 5 4 4 4 2 3 2 2 3 49153 D A T U MD A T U MD A T U M3 2 2 3 3 4 3 3 3 3 3 2 3 2 4 3 4 47To further refine the selected design, a three-dimensional graphical simulation software was used to identify the suitable working volume for the system for two different fabrication situations; (i) simple fabrication with no part rotation and tilting and system is required to do vertical buildup only, (ii) more complex fabrication requiring part rotation and tilting and the system is depositing in the vertical direction.

For both conditions, work perimeter of 12” x 12” x 12” and 8”x 8”x 8” were simulated. CNC codes were written to instruct the system to perform tool movement to cut the peripheral of the specified work perimeter, and any obstruction or collision was recorded. Simple movement is represented by a cube, while a more complex movement is represented by a part with overhang as illustrated in figure 4(a) and figure 4(b).(a) Cube representing simple vertical buildup (b) Part requires rotation and tiltingFigure 4: Representative objects for tool path and work volume verificationFrom the simulation results, the 12” x 12” x 12” work perimeter had demonstrated some collisions for rotating and tilting movement, while the 8” x 8” x 8” work perimeter had shown no collision for both simple and rotate-tilt movements. The results are presented in Table 5.

Major hardware components of the system is shown in Table 6Table 5: Simulation Results for 12” x 12” x 12” and 8” x 8” x 8” work volumeWork Envelope RepresentationsDeposition ProcessMilling ProcessPerimeter (12” X 12” X 12”)No collisionNo collisionCollision between part andWith rotation and tilt (12” X 12” X 12”)Collision between part and rotary table baseCollision between part and rotary table baseCollision between part and machine column.Perimeter (8” X 8” X 8”)No collisionNo collisionWith rotation and tilt (8” X 8” X 8”)No collisionNo collision16Table 6: Major Hardware Components of LBMD SystemComponentSub-ComponentsPrimary Functions2-Axis Rotary TableRotation and tilt motion • Work table for the system5-axis CNC Machine3-Axis milling machineX, Y and Z motion • Controls all axes motionPowder Storage HopperStore metal powder • Releases powder to powder feeder unitPowder Feeder UnitFeeds powder into nozzle • Controls powder flow ratePowder and Laser Delivery SystemPowder and Laser NozzleDeliver powder to melt pool • Deliver protection and carriage gas • Holds laser focus lensesLaser HeadGas Supply SystemProvide Protection and Carriage gasFiber Optic CableDeliver laser beam to focus lensesWay coverProtect Linear table from contaminants (dust and moisture)1-axis Linear TableMotor and EncoderProvide Linear table motion • Closed loop control for linear table4 SOFTWARE AND CONTROL SYSTEMTypical software architecture for a Laser Based Material Deposition includes the modeling of a three-dimensional CAD model in standard Stereolithography (STL) format, generation of layer representation of the object which is equal to the deposition thickness, and creation of CNC codes for the tool path that are understandable by the machine controller [3,4,7].

The system architecture for the LBMD system is illustrated in Figure 5.17LBMDDesignMass Flow RateCAD DataPower InformationPowder Feeder ControllerCutting path, Feed rate and Tool informationLaser ControllerPowder FeederCNC ControllerPowderSplitterLaserPowerNozzlePowderCNC MillingProcess MonitorProductFigure 5: System Architecture [8].5 CONCLUSIONSDesign and integration process of a Laser Based Material Deposition System have been presented and a LBMD system is proposed. A three-dimensional graphical simulation has been applied and proven to be a useful tool in the development of the LBMD system. A few important considerations in the design and integration of the system include the following: (i) system performance – high accuracy and precision, stable powder and power delivery, online control and monitoring. (ii) high temperature working environment – nozzle made of high temperature material and equipped with cooling system.

(ii) high temperature working environment – nozzle made of high temperature material and equipped with cooling system. (iii) protection of optical lenses or any sensitive components from moisture and dust or easily cleaned and replaceable. (iv) safety – laser radiation, metal powder, coolant mist and dust. (v) maintenance – laser alignment check should be minimize, ventilation system is needed to properly disposed contaminants (metal powder, coolant mist, dust) out of the systemACKNOWLEDGEMENTSThe support from the National Science Foundation Grant No. DMI 9871185 is appreciated. The contributions from the RP group at UMR and the support from Research and Development Unit, Universiti Sains Malaysia is also noted here.18REFERENCES1. Jacobs, P.F., Rapid Prototyping & Manufacturing: Fundamentals of Stereolithography, Society of Manufacturing Engineer, 1st edition, pp.1, 18-20, 1992.2. Koch, J.L.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesign and Manufacture of a Continuous Fiber-Reinforced 3D Printed Unmanned Aerial Vehicle WingDhileep Kumar Jayashankar1, Aarthi Devarajan1, Guoying Dong2, David Rosen31Research Assistant, 2Research Fellow, 3Professor Digital Manufacturing and Design (DManD) Centre, Engineering Product Development Singapore University of Technology and Design, SingaporeAbstractThe Markforged Mark Two 3D printer is capable of printing various orientations of continuous fiber reinforcement. An initial study of how the orientation of the fiber influences the strength characteristics (tensile and flexural properties) was conducted.

An initial study of how the orientation of the fiber influences the strength characteristics (tensile and flexural properties) was conducted. Four combinations of carbon fiber reinforcement orientations were tested, specifically unidirectional, isotropic, concentric and a combination of isotropic and concentric, with the Markforged Onyx matrix material. The results will aid in designing a wing with the optimum fiber configuration that will give the desired mechanical properties based on the forces acting on the wing. Design for Additive Manufacturing (DfAM) concepts and tools will be used to design and manufacture a large UAV wing. Topology optimization, based on a CFD computed pressure distribution, was used to determine geometric regions where carbon fiber reinforcement could be best utilized. From there, a honeycomb structure was designed to ensure stiffness and light weight based on desired densities.

From there, a honeycomb structure was designed to ensure stiffness and light weight based on desired densities. A wing section was fabricated using the Mark Two printer to identify the capabilities and limitations of the system in realizing the design objectives.Keywords: UAV wing, topology optimization, fiber reinforced composites, material extrusion, design for additive manufacturing1IntroductionAmong the various additive manufacturing (AM) processes, the material extrusion (MEX) process has the capability to produce functional parts in a variety of engineering thermoplastics for various applications [1]. In the past decade, 3D printing of composite materials through MEX had received significant attention among researchers. Composite 3D printing is a process of combining multiple materials in an established 3D printing process to improve mechanical and thermal properties such as strength, stiffness, durability and heat resistance[2].

In AM, composites are based on mixtures of matrix materials and reinforcements. Our interest in this work is in fiber reinforced thermoplastics with the matrix materials including ABS, nylon, polycarbonate, and PLA. Thermoplastics possess specific properties like malleability, ductility and low melting point, which make them an excellent fit for AM. Fiber reinforcements, such as continuous carbon fiber, significantly increase strength and stiffness which is important in a variety of applications. Options for reinforcements include short fibers and continuous fibers which could be carbon, Kevlar or glass fiber. Continuous fibers have the ability to form large structural elements that can withstand substantial loads and disseminate forces along the directions of fiber[3].

Continuous fibers have the ability to form large structural elements that can withstand substantial loads and disseminate forces along the directions of fiber[3]. Ultimately, this property780of fibers provides advantages in strength over traditional thermoplastics, which makes them a promising means of reinforcement in AM to build light-weight parts [4].However, the manufacturing compatibility of fiber printing is complicated and poses a significant challenge when it comes to mixing the reinforcement with polymer matrix due to control of fiber orientation, void content and low-cost production [5]. Some researchers have identified three possibilities for fiber integration into the part in MEX processes, which are: 1) adding fibers before the printing process, i.e., filaments itself are a composite; 2) adding fibers in the print head, i.e. one extruder with two entries for filaments; and 3) adding fibers in the part, i.e.

one extruder with two entries for filaments; and 3) adding fibers in the part, i.e. with two separate extruders for materials (one for matrix and another for reinforcement) [6]. Approach 1) is useful for carbon nanotubes or graphene reinforcements [7]. Our interest is in approach 3) where we use the Markforged Mark Two printer with their Onyx material and continuous carbon fibers. The Onyx material has a nylon matrix and short carbon fibers. Continuous fibers are deposited by a second printhead that can selectively reinforce each layer.The on-going research in 3D printing (Type 3) is based on how effectively fibers can reinforce thermoplastics to increase the mechanical properties, which play a major role in advancements in both academic research and for commercial products. The most common matrix-reinforcement combinations in fiber printing are Nylon or ABS-carbon fiber, PC-carbon nanotubes, and PLA- graphene [7].

The most common matrix-reinforcement combinations in fiber printing are Nylon or ABS-carbon fiber, PC-carbon nanotubes, and PLA- graphene [7]. As th detailed evaluation of mechanical properties of additively manufactured Nylon-carbon fiber reinforced composites are very limited to the authors' knowledge, in this work, we attempted to use Markforged Mark II 3D printer to find the mechanical properties of various Onyx-carbon fiber combinations and applied that specific material combination to 3D print a UAV wing structure.Design for Additive Manufacturing (DfAM) is another paradigm that enables the designers to maximize the product performance through the synthesis of shapes, sizes, hierarchical structures, and material compositions, subject to the capabilities of AM technologies[8]. By implementing process-specific design rules given by part-level DfAM, one can create the product-process design space and achieve the functional optimization of the design[9].

By implementing process-specific design rules given by part-level DfAM, one can create the product-process design space and achieve the functional optimization of the design[9]. On the other hand, research on Assembly-level DfAM is still at initial phase to redesign the assemblies by taking advantage of AM-enabled capabilities and its design freedom[10]. In assembly-level DfAM, the possibilities of 3D printing functional assemblies, designing novel product architectures, and achieving part consolidation remain under investigation. We applied in designing a lightweight UAV wing structure using DfAM principles.In this work, we investigated part-level and MEX-specific DfAM rules for designing a lightweight wing structure of a UAV. We explored a sequential product-process design optimization method and implemented the concepts in designing the UAV wing structure.

We explored a sequential product-process design optimization method and implemented the concepts in designing the UAV wing structure. In support of this design project, the tensile and flexural mechanical properties of 3D printed continuous carbon fiber composites were characterized for four different fiber infill configurations. Resulting properties, as well as the capabilities of the Markforged Mark Two system, were utilized in the wing design. At last, this article also discusses about the manufacturing challenges and limitations of fiber reinforced 3D printing.7812 Materials and methods2.1 Materials The materials, such as Onyx (micro carbon fiber filled nylon/mixture of nylon with chopped carbon fiber) and continuous carbon fiber filament (CCF) were purchased from Markforged, United States. The materials had been conditioned in a dry cabinet with humidity level at 25% prior to use to avoid moisture entrapment.

The materials had been conditioned in a dry cabinet with humidity level at 25% prior to use to avoid moisture entrapment. The diameter of the Onyx filament and CCF fiber were measured to be 1.7 mm and 0.32 mm. And the density of Onyx and CCF were measured to be 1.2 g/cm3 and 1.4 g/cm3 correspondingly.2.2 Manufacturing The designed parts were fabricated using Markforged 3D printer with different combination of carbon fiber and Onyx matrix. Of the different CF printers available in today’s market, Mark One, Two and X7 from Markforged are particularly relevant for a wide range of applications including aerospace, and automobile, etc. In this study, we used the Mark Two 3D printer at the Singapore University of Technology and Design to print the parts as shown in figure 1(a). In the composites world, Markforged coined this fiber printing technology as Continuous Filament Fabrication (CFF), which comes under the category of MEX processes.

In the composites world, Markforged coined this fiber printing technology as Continuous Filament Fabrication (CFF), which comes under the category of MEX processes. These printers utilize two nozzles to print the part, one being a thermoplastic nozzle with Onyx (nylon matrix with chopped carbon fiber that also acts as a support material), and another one for carbon, glass or kevlar fiber. The Mark Two printer has a provision to mount a CF filament spool and a separate pelican box for matrix material.Figure 1 (a) Mark II 3D printer (b) Flexure specimen sliced in EIGER softwareThe volume fraction of fiber to matrix can be easily controlled by the pre-processing software called EIGER, which is a cloud-based software provided by Markforged that can be used to import, orient and align the geometrical parts in the print bed before printing.

The user-machine interface provides the user with two customized fill options, concentric and isotropic infill, in which the user can control the fiber deposition at each layer in a part using customizable fiber angle and orientation options. Within this paradigm, one can combine both concentric & isotropic infill and generate several different fiber configurations in a part based on the performance requirements and cost. In our case, the .STL files were imported in the EIGER software to prepare parts for printing by inducing the fibers at required locations as shown in figure 1(b). The samples were then printed with the predetermined print settings provided by Markforged as there is lack of option for customized print parameters. Throughout the printing, the Onyx nozzle temperature was maintained at 274°C and CCF’s nozzle temperature at 254°C.

Throughout the printing, the Onyx nozzle temperature was maintained at 274°C and CCF’s nozzle temperature at 254°C. The brim option was enabled in the print settings, and washable glue was applied to the bed plate for part-bed adhesion.7822.3 Design of fiber configurations The orientation of the carbon fiber plays a major role in strength characteristics for both tensile and flexural properties. With the available fiber fill options, we generated four possible combinations of fiber reinforcement by combining both concentric and isotropic fill along with Onyx as shown on Table 1.Table 1 Fiber Reinforcement ConfigurationsAccording to ASTM standards, the sample thickness for tensile test is set be 2.5 mm with 20 layers of printing, out of which 16 layers are utilized for fiber reinforcement. Similarly, the sample thickness for flexural test is set to be 4 mm with 32 layers of printing, out of which 28 layers are utilized for fiber reinforcement.

Similarly, the sample thickness for flexural test is set to be 4 mm with 32 layers of printing, out of which 28 layers are utilized for fiber reinforcement. For the sake of simplification, we represent the samples using codes throughout the article. The naming conventions for the samples are Tensile (T), Flexural (F), Unidirectional (U), Isotropic (I), Concentric (C), and Isotropic & Concentric (I_C). The sample codes were generated based on the type of test method used followed by the fiber configuration, for example, T_U represents tensile samples with unidirectional fiber layers. The first configuration, F_U & T_U was designed with unidirectional fiber reinforcement, the fibers being oriented at 0° angle at all fiber layers. The second configuration, F_I & T_I was designed with isotropic fiber reinforcement, the fibers being oriented at 0°, 45°, 90°, and 135° angles, where each angle takes up an entire layer in the sample.

With the isotropic fiber configuration, once the base matrix layers are printed, the first fiber layer would be laid at [0°] angle followed by [45°], [90°] and [135°] at second, third and fourth layers, respectively. The fiber angle pattern [0/45/90/135°] is then repeated at every four layers which in our case is considered as the Isotropic (I) configuration. The third configuration, F_C & T_C was designed with concentric fiber reinforcement, four fiber rings [0/90°] being concentrically surrounded from outwards to inwards without isotropic fill. The fourth configuration, F_I_C & T_I_C was designed with both types of783fiber reinforcement, with isotropic fibers oriented at [0/45/90/135°] angles in the center of the specimen followed two concentric fiber rings [0/90°] surrounding them.The fiber volume fraction for all the configurations was estimated from the Eiger-cloud slicing software provided by Markforged.

The calculation is akin to traditional composite structures that follows the rule of mixtures (ROM) method. Since, Eiger software has the ability to compute the required volume of both matrix and fiber material for the part to be printed, the fiber volume fraction was estimated from the software itself and reported at figure 2. However, the experimental study was not conducted for the distribution of voids and porosity.Figure 2 Fiber volume fraction for different fiber configuration2.4 Experimental 2.4.1 Tensile test The test specimens were designed and printed according to ASTM D3039 (Standard Test Methods for Tensile Properties of Polymer Matrix Composite Materials) as shown on figure 2. The dimensions of the specimen were set to be 25 × 250 × 2.5 mm. Five specimens were 3D printed for each type of fiber configuration. The test was conducted in a double column universal testing machine - 100 kN Instron 5982 (Instron, US).

Five specimens were 3D printed for each type of fiber configuration. The test was conducted in a double column universal testing machine - 100 kN Instron 5982 (Instron, US). The speed of testing was calculated from the ASTM standards and set to be 1.5 mm/min.2.4.2 Flexural testThe test specimens were designed and printed according to ASTM D7264 (Standard Test Methods for Flexural Properties of Polymer Matrix Composite Materials) as shown on figure 3. Considering the span-to-depth ratio as 32:1 and support span length as 128 mm, the dimensions of the specimen were set to be 13 × 168 × 4 mm. Five specimens were 3D printed for each type of fiber configuration. The test was conducted in a single column universal testing machine - 1 kN Instron7845943 (Instron, US).

Five specimens were 3D printed for each type of fiber configuration. The test was conducted in a single column universal testing machine - 1 kN Instron7845943 (Instron, US). The speed of testing was calculated from the ASTM standards and set to be 6.82 mm/min.Figure 3 Dimensions of tensile specimen and sectional view to illustrate the fiber and matrix layersFigure 4 Dimensions of flexural specimen and sectional view to illustrate the fiber and matrix layers2.4.3 Tensile PropertiesFigure 4(a-d) shows the typical tensile stress vs strain graph of five specimens in four designed fiber configurations, such as (a) Unidirectional (T_U), (b) Isotropic (T_I), (c) Concentric(T_C) and (d) Isotropic & Concentric (T_I_C). It is observed that the unidirectional fiber configuration specimens (T_U) exhibited higher tensile stress and modulus of elasticity than other fiber configurations with 523.98 ± 20.1 MPa and 23.59 ± 0.976 GPa, respectively.

Similar results were reported by Goh et.al [11] but all T_U coupons broke abruptly due to the brittle nature of the carbon fiber without any drops which contradicts the statement reported by Goh et.al.[11] The rest of the fiber configurations, T_I, T_C, and T_I_C behaved in a very similar way but with strength and stiffness values nearly half of those for T_U specimens with the range between 180-230 MPa for maximum tensile strength and 10-14 GPa for modulus of elasticity. T_U & T_I specimens exhibited brittle failure and the fracture surface appeared to be fibrous husk where the fibers spread out shortly after the breakage. With brittle failure in T_U & T_C specimens, the fiber-matrix785debonding was the primary fracture mechanism as the matrix material was separated from the fiber like a shell as shown on figure 5(b).

The flexural test results demonstrate that the strength and stiffness of the specimens were greatly affected by the variation of fiber reinforcement and the fiber volume fraction as shown on figure 2. From figure 6(a), it is observed that unidirectional fiber configuration specimens (F_U) exhibited highest flexural strength and modulus than other fiber configurations with 352.51 ± 14.01 MPa and 22.97 ± 0.15 GPa, respectively. F_U specimens behaved linearly up to 1.4% flexural strain and failed abruptly as shown on figure 7(b) which is similar to the studies reported by Goh et.al [11]. On the other hand, F_C & F_I_C specimens behaved analogous to F_U sample until reaching 1.5% strain but with F_I_C configuration, it was noted that the samples were not broken even after excessive straining time.

Both F_C & F_I_C designs appeared to show successive fiber failure with plastic yielding and based on the stress- strain graph, it is clear that the specimens suffered fiber micro-cracking after 2% strain is achieved. Also, the specimens with high fiber volume fraction achieved significantly higher strength and show low ductility. Surprisingly, the concentric fiber configuration (F_C) performed well by having only 4 fiber rings with the very low fiber volume fraction of 13.28% as shown in Figure 2.

Since the MQ-9 design is one of the most successful UAV in the last two decades, it was chosen for a better comparison for the proposed design techniques.Some parts of the design were acquired from several online databases and simplified according to our study as our focus is mainly on optimizing the wing structure. Only one side of the UAV wing structure is considered for this study and to 3D print the full wing structure in the Mark Two, the half-wing was scaled down to 16:1 with the span of 625 mm and the root and tip chord as 90 mm & 45 mm.

For the sake of simplicity and in order to demonstrate the design methodology for wing structure involving additive manufacturing, the acquired wing design will act as a supplement for validating the design method.3.2 Numerical modelling of the UAV wing structure The ultimate aim of this study is to conduct topology optimization on a wing structure to reduce the mass while maintaining the structural performance and taking advantage of design freedom through additive manufacturing.

Numerical studies including computational fluid dynamics (CFD), finite element analysis (FEA) and topology optimization were conducted on the half wing structure of the UAV using SolidWorks 2020 software.788Figure 9 Design of Unmanned Aerial Vehicle (Predator B - MQ-9 Reaper) with a, b, c being top view, front view, and Isotropic view(a) Half wing design from UAV(b) Mesh model of the half wing for CFD & FEAAnalysisFigure 10 Half wing structure from UAV designThe analysis begins with the CFD of the wing structure and the results from CFD were taken as input loading conditions in the static analysis to calculate von-Mises stresses and displacements, which was then followed by topology optimization. As the velocity of most of the UAV falls between 45-60 m/s, we chose 50 m/s for CFD study, then the results from CFD study in the form of pressure distribution on the wing structure and the gravitational force were considered for the static FEA.

At last, the topology optimization study involves the result from both CFD and FEA and the weight optimization parameters were set to 5%, 10% and 15% of the weight of the solid wing structure. The FEA component in SolidWorks was used for the analysis. For the initial studies, only Onyx material was used for the entire wing structure whereas in the789later part Onyx and continuous carbon fiber were used for redesigning the wing. Figure 10(a & b) shows the half wing design from UAV and its SolidWorks mesh model.4 Result & Discussion4.1 Stresses and Finite Element Analysis The primary objective of the CFD analysis is to find out the dynamic pressure distribution over the wing structure through an input velocity of 50 m/s in x-axis (towards the leading edge). Within SolidWorks, under the CFD domain, the density of air was set to be 1.225 kg/m3 and the thermodynamic parameters such as pressure and temperature were set be 101325 Pa and 293.2 K.

Within SolidWorks, under the CFD domain, the density of air was set to be 1.225 kg/m3 and the thermodynamic parameters such as pressure and temperature were set be 101325 Pa and 293.2 K. Furthermore, the thermal wall condition for the analysis is chosen as adiabatic wall with 5 µm roughness. Figure 11 shows the CFD results and the pressure distribution over the wing surfaces. As the wing already experiences dynamic loads throughout the leading edge, it is evident that the leading edge needs to be strengthened by introducing carbon fiber in the design. However, no stress and load were spotted at the trailing edge of the wing. From the CFD results, the pressure distribution in 3D space was exported and given as an input loading condition for the static analysis.Figure 11 Static loading conditions obtained from computational fluid dynamics (CFD) analysisFigure 12(a & b) shows the von-Mises stress and displacement of the wing from the static analysis.

The output of the static study was given as an input for the topology optimization and the wing is optimized for best stiffness to weight ratio for three different weight percentages, 5%, 10% and 15%, as shown on figure 13.790(a) von-Mises stress(b) DisplacementFigure 12 Static analysis of the wing(b) Weight reduction to 10%(a) Weight reduction to 5%(c) Weight reduction to 15%Figure 13 Topology optimization for different weight percentagesThough the mechanical properties of 3D printed carbon fiber composites has been analyzed and evaluated, there are other considerations such as the effect of print process parameters that includes fiber orientation, fiber volume fraction, stacking sequence of fiber ply, matrix infill pattern (such as triangular, rectangular, hexagonal, etc), infill density, raster angle, layer thickness, build orientation (XY, YZ, ZX), printing temperatures (Onyx nozzle temperature, fiber nozzle temperature), and printing speed.

Since a process parameter study requires extensive research on fiber composites, it was not included in this study.4.2 Optimized UAV wing structure The optimized wing structure was designed in SolidWorks and Rhino software based on the resultant geometric configuration from the topology optimization in which the weight of the wing was reduced to 5% of its original weight. The optimized design was exported and remodeled to incorporate honeycomb design and carbon fiber reinforcement in the wing. Since the final part is subjected to 3D printing, the wing was designed in a way that is compatible to be produced by the791Mark Two machine. The new wing design consists of three parts; namely, (i) Outer shell; (ii) Carbon fiber reinforcement; and (iii) Honeycomb core. Figure 15(a) & (b) offers a closer look of the final wing design with 5% density design in which the outer shell was set to be transparent for the visibility of core and fiber.

Figure 15(a) & (b) offers a closer look of the final wing design with 5% density design in which the outer shell was set to be transparent for the visibility of core and fiber. The fiber was reinforced in the top and bottom sections of the wing that resulted from topology optimization for each of the three different weight percentages shown on figure 13. Additionally, we placed the fibers at the leading edge of the wing as it undergoes high stress as obtained by CFD and static analysis. In contrast, the stress at the trailing edge was comparatively smaller so it was ignored. The carbon fibers were bolstered in the design at 0° orientation as the unidirectional (T_U & F_U) fiber placement attained higher strength than other fiber configuration as discussed in section 3.1 and 3.2.The 2D regular honeycomb core was designed based on the relative density model from cellular solids [12].

For designing the honeycomb core, we developed a computational parametric model in Rhino-Grasshopper software that can compute the side length of the hexagon (l) and cell wall thickness (t) as shown on figure 14(a) while still maintaining its relative density to 5%. Using the relative density equation (as mentioned below), we found out the ratio of cell wall thickness (t) to side length of hexagon (l) as 0.0433 by manual calculation. Then, t/l ratio was given as an input parameter in Rhino-Grasshopper software as shown on figure 14(a) to find out the honeycomb’s cell size and cell wall thickness that satisfies the above relation. Once the raw honeycomb is generated in Rhino-Grasshopper, the interior wing profile was designed by leaving sufficient space for both fibers and the outer shell. The interior profile was then imported into Rhino and the actual honeycomb core was cut from the raw honeycomb section using some basic Boolean operations.

The interior profile was then imported into Rhino and the actual honeycomb core was cut from the raw honeycomb section using some basic Boolean operations. At last, both the honeycomb core and the outer shell (with the space for fiber reinforcement) was assembled in SolidWorks as shown on figure 14(c).(a) Grasshopper setup to generate honeycomb structure that satisfies t/l ratio(b) Cell shape and labelling of regular honeycomb;(c) Exploded assembly generated honeycomb and shell design inRhino-Grasshopper softwareFigure 14 Optimized UAV wing design process792Relative Density,𝜌∗ 𝜌𝑠=𝑀𝑠 𝑉𝑇𝑉𝑠 𝑀𝑠=𝑉𝑠 𝑉𝑇= 0.05𝑊ℎ𝑒𝑟𝑒,𝜌∗ = 𝑑𝑒𝑛𝑠𝑖𝑡𝑦 𝑜𝑓 𝑐𝑒𝑙𝑙𝑢𝑙𝑎𝑟 𝑠𝑡𝑟𝑢𝑐𝑡𝑢𝑟𝑒(𝑣𝑜𝑙𝑢𝑚𝑒 𝑓𝑟𝑎𝑐𝑡𝑖𝑜𝑛 𝑜𝑓 𝑠𝑜𝑙𝑖𝑑)𝜌𝑠 = 𝑑𝑒𝑛𝑠𝑖𝑡𝑦 𝑜𝑓 𝑐𝑒𝑙𝑙𝑢𝑙𝑎𝑟 𝑠𝑜𝑙𝑖𝑑𝜌∗ 𝜌𝑠=2√3𝑡 𝑙= 0.05 ⇒𝑡 𝑙= 0.0433𝑀𝑠 = 𝑚𝑎𝑠𝑠 𝑜𝑓 𝑐𝑒𝑙𝑙𝑢𝑙𝑎𝑟 𝑠𝑡𝑟𝑢𝑐𝑡𝑢𝑟𝑒𝑉𝑇 = 𝑣𝑜𝑙 𝑜𝑓 𝑐𝑒𝑙𝑙𝑢𝑙𝑎𝑟 𝑠𝑡𝑟𝑢𝑐𝑡𝑢𝑟𝑒(𝑓𝑜𝑟 𝑎 𝑟𝑒𝑔𝑢𝑙𝑎𝑟 ℎ𝑒𝑥𝑎𝑔𝑜𝑛)𝑉𝑠 = 𝑣𝑜𝑙 𝑜𝑓 𝑠𝑜𝑙𝑖𝑑 (𝑖. 𝑒.

𝑒. 𝑤𝑖𝑡ℎ𝑜𝑢𝑡 𝑣𝑜𝑖𝑑𝑠)(a) Optimized wing and assigned the(b) Optimized wing structure with honeycomb design and carbonmaterial for 3D printing in Markforgedfiber reinforcementFigure 15 Optimized UAV wing design(a) Small section of final wing design sliced using Eigercloud slicing software(b) 3D printed model in Mark II – the top outer skin of the wing was removed for visibilityFigure 16 Optimized UAV wing sliced in EIGER, followed by printing in Markforged Mark IIA small section of the wing model was created and imported in build preparation software- Eiger for visualizing fiber placements and configuring the build for 3D printing.

While designing the final wing, we introduced the fiber at the leading edge considering the stresses involved in CFD and static analysis but since the cross section of the fiber at leading edge is relatively small, and possesses a tapered section, the build preparation software failed to fill the fiber in that region793as shown in Figure 16 (a). It became evident that the Markforged system has some limitations that were not known to us. Such issues require users to do more iterations in designing fiber infills. The small section was successfully 3D printed as shown in Figure 16 (b); the top skin was removed so that the internal structures are visible.

The small section was successfully 3D printed as shown in Figure 16 (b); the top skin was removed so that the internal structures are visible. In the future, we will use our learning about the Markforged system to redesign the fiber reinforcement to properly account for the system’s capabilities and limitations.5 ConclusionThe Markforged Mark Two 3D printer is a type of material extrusion system that is capable of printing various orientations of continuous fiber reinforcement. We designed a scale model of a UAV wing to be fabricated using the Markforged system. In support of this design study, we characterized the mechanical properties of Markforged fabricated continuous carbon fiber reinforced specimens. Four combinations of carbon fiber reinforcement orientations were tested, specifically unidirectional, isotropic, concentric and a combination of isotropic and concentric, with the Markforged Onyx matrix material. Both tensile and flexural properties were measured.

Both tensile and flexural properties were measured. To design the UAV wing, a configuration including topology optimized stiffeners with a honeycomb structure was chosen. The wing skin and honeycomb were to be fabricated with the Markforged Onyx material without continuous carbon fiber reinforcements, while the topology optimized sections had continuous fibers. A wing section was fabricated to test the design.Based on the project to date, the following conclusions were drawn:The unidirectional continuous fiber reinforcement orientation was the strongest and stiffest, by far, of the tested orientations. All tested orientations exhibited brittle fracture in tension testing, while the isotropic and concentric combination exhibited the most ductile failure in flexural testing.Although the concentric orientation contained only 4 fiber rings and the lowest fiber volume fraction, it had the second best flexural properties.

The mechanical property results indicate that mechanical properties of fabricated parts will be highly sensitive to local fiber orientation. • Wing designs with honeycomb reinforcements are well suited to the Markforged and, more generally, MEX processes. With optimized fiber reinforced regions, resulting wing designs can possess high stiffness to weight ratios.The Markforged system makes certain assumptions about depositing fiber reinforcements based on local geometry size and shape that are not necessarily evident to users of the system. Experience is needed to anticipate issues when designing small or narrow regions in which fiber reinforcements are desired. AcknowledgementThe authors gratefully acknowledge support from the Singapore DSO under grant DOSCL20019. Support from the Digital Manufacturing and Design Centre at SUTD, funded by the Singapore National Research Foundation, is also acknowledged for access to the Markforged printer.References[1]I. Gibson, D.W.

SIGN AND MANUFACTURE OF A FORMULA SAE INTAKE SYSTEM USING FUSED DEPOSITION MODELING AND FIBER-REINFORCED COMPOSITE MATERIALSRyan Ilardo, Christopher B. Williams Design, Research, and Education for Additive Manufacturing Systems Laboratory Department of Mechanical Engineering Virginia Polytechnic Institute and State University Randolph Hall, Blacksburg, VA, 24061-0238 (540) 231-3422; cbwilliams@vt.eduABSTRACTIn this paper, the authors discuss the design and manufacture of an intake system for a 600cc Formula SAE engine. Specifically, Fused Deposition Modeling is used to create an intake system (consisting of a plenum, plenum elbow, and cylinder runners) that is then later covered in layers of carbon fiber composite fabric through vacuum bagging.

As a result of this approach, the geometry of the intake system has been redesigned to result in reduced weight (due to lower material density and lack of welds, hose clamps, and silicon couples), improved charge distribution, and increased torque through a wide RPM range when compared to its traditionally-manufactured aluminum counterpart.Keywords: Fused Deposition Modeling, Intake Manifold, Formula SAE1. CONTEXT: INTAKE MANIFOLD OF A FORMULA SAE VEHICLE1.1 Formula SAEFormula SAE is an international student design competition organized by the Society of Automotive Engineers (SAE). In this competition, student design teams design, build, and test a small Formula-style race car. The teams, acting as contractors for a fictional manufacturing company, are held to strict rules and design the car as if it were a production item. The car is evaluated via static and dynamic tests.

The teams, acting as contractors for a fictional manufacturing company, are held to strict rules and design the car as if it were a production item. The car is evaluated via static and dynamic tests. The static evaluations are comprised of an analysis of the car’s cost and design as well as a team presentation. The dynamic testing events include skid pad, acceleration, autocross, and endurance trials.In an effort to create a fuel-efficient, lightweight, cost-effective, and fast automobile, students carefully design each component of their vehicle. Subsystems, ranging from the vehicle frame, to the steering system and the powertrain, are designed by the students, and are frequently manufactured in- house.

Subsystems, ranging from the vehicle frame, to the steering system and the powertrain, are designed by the students, and are frequently manufactured in- house. In this paper, the authors focus on the design and (additive) manufacture of the intake manifold system of the 2009 Virginia Tech Formula SAE car.1.2 Intake ManifoldWhen designing the engine package for a Formula SAE car, as well as other automotive applications, it is very important to design a quality intake system. The primary function of the intake manifold system (Figure 1) is to deliver combustion air to the engine. Specifically, the primary design goal is to distribute the combustion mixture evenly to each intake port, as doing so improves the engine’s ability to efficiently and effectively produce torque and power.

The geometric design of the intake system affects the volumetric efficiency of the engine, and thus directly affects the performance of the vehicle.To realize the primary design goal of providing equivalent amounts of air to each cylinder, there areseveral objectives to consider when designing an intake manifold:Minimize pressure loss, as pressure loss results in a decrease in output power. • Maintain equal static pressure distribution in the plenum, as this will cause the cylinders to pull the same vacuum, thus leading to even flow in each cylinder.Minimize bends and sudden changes in geometry, as these geometric affects can cause pressure loss.Maximize air velocity into the cylinder, as this provides a better mixture of fuel and air, which results in better combustion and performance.Minimize the mass of the system; a common goal of every subsystem of the vehicle.PlenumPlenum ElbowCylinder RunnerFigure 1.

3D rendering of the complete intake assembly.1.3 Intake Manifold ManufactureTraditionally, intake manifold systems are manufactured via bonding multiple aluminum pieces (typically either bent pieces of aluminum tubing or casted aluminum components) together with multiple welds. An example of the aluminum intake designed and manufactured by the 2008 Virginia Tech Formula SAE team is shown in Figure 2. While the use of aluminum provides a structurally sound component with a low mass, this manufacturing technique fundamentally limits the types of manifold geometries that can be created. Due to the inherent geometric limitations imposed by the existing manufacturing process (bending and welding), the team found it difficult to design and fabricate a system in which (i) pressure losses were kept to a minimum and (ii) equal charge was provided to each cylinder.In this paper a new manufacturing process for fabricating intake manifolds for Formula SAE race cars is presented.

Specifically, an intake manifold system is fabricated in a two-step process: (i) a thin-walled ABS core is created via Fused Deposition Modeling (FDM) and then (ii) wrapped with layers of carbon fiber via vacuum bag curing. The combined use of additive manufacturing and composite materials allows for the manufacture of a functional intake with virtually any geometry. This geometric freedom offered by FDM allows a designer to ensure even static pressure distribution throughout the intake and an equal charge to each cylinder. Furthermore, as FDM allows the creation of the entire manifold system as a single piece, the need for welds and other fasteners is eliminated, resulting in an expected reduction in mass.An overview of the proposed manufacturing process is presented in Section 2. The design and manufacture of an intake manifold using the process is detailed in Section 3.

The design and manufacture of an intake manifold using the process is detailed in Section 3. In Section 4, the authors present performance results of the manifold and compare them with manifolds created by more traditional means. Finally, closure is offered in Section 5.Figure 2. Intake manifold composed of welded aluminum pieces.2. ADDITIVE MANUFACTURE OF AN INTAKE MANIFOLD2.1 Prior EffortsRecognizing the geometric flexibility offered by additive manufacturing (AM) techniques, engineers have looked to the technologies as a means of improving the design of intake manifold systems for high- performance automobiles. Various AM technologies have been used to create functional prototypes for testing, to fabricate tooling for creating end-use parts, or to create the end-use parts themselves.Engineers at Jaguar Cars Ltd. have used laser sintering to prototype intake manifolds for functional testing.

have used laser sintering to prototype intake manifolds for functional testing. The use of these prototyping techniques increased the amount of design iterations while decreasing the financial investment and overall development time for this component [1]. Other engineers have reported the use of laser sintering [2] and solid ground curing [3] to create intake manifold prototypes suitable for testing.Additionally, AM techniques have been used to create tooling for the manufacture of end-use intake manifolds. In these processes, a pattern of the manifold is created via an AM technology and is then transformed into a functional component either through coating with a carbon fiber composite (followed by removal/dissolution of the pattern) [4,5], or through coating with a ceramic for investment casting with aluminum [6].Finally, many engineers have looked to AM techniques as a means of creating end-use, functional manifolds.

Specifically, several SAE teams from around the globe (Australia [7], United Kingdom [8, 9], United States [10], and Greece [11]) have used laser sintering (LS) to create functional, race-ready manifolds from sintered nylon. To the authors’ knowledge, laser sintered nylon is the only reported AM technique that has been used to create an end-use manifold. Nylon is an appropriate material for this application as it is robust enough to handle the high operating temperatures and pressures of the intake manifold.2.2 Fused Deposition Modeling + CompositesLacking access to a LS machine (and having insufficient funds to hire a service bureau), the student design team looked to other AM technologies as a means of creating an end-use manifold. Having access only to FDM technology (a Stratasys Dimension SST), the students were aware that it alone would not be capable of producing an end-use part.

Having access only to FDM technology (a Stratasys Dimension SST), the students were aware that it alone would not be capable of producing an end-use part. The standard ABS material (Stratays, P400 [12]) is neither strong enough to withstand the high pressures found in a turbocharged engine, nor it is heat-resistant enough to its high temperatures. While Stratasys does provide high-temperature and high-strength material options (PPSF/PPSU [13], ULTEM 9085 [14]), they can only be processed on Stratays’s higher-end machine line, Fortus. Furthermore, parts created via FDM are not airtight due to pores created by poor optimization of material flow, filament/roller slippage, liquefier head motion, and build/fill strategies in the extrusion deposition process [15].In an effort to address these processing deficiencies, the student team proposed a manufacturing process in which a thin-shelled, FDM-fabricated part is layered with composite material.

Specifically, the team chose to apply carbon fiber fabric to the ABS part with a high-temperature resin via a vacuum bagging process. This process alleviates the geometric limitations of traditional manifold manufacturing techniques, and also ensures a light-weight, strong, and heat-resistant component. Additional benefits of this process include:While an additional post-processing step is needed in this process when compared to direct fabrication using LS (Section 2.1), the equipment required to fabricate the manifold in the proposed process is much less expensive; LS machines typically cost more than an order of magnitude greater than FDM machines.Thanks to the added strength provided by the carbon fiber composite material, the printed manifold shell can be very thin (0.12” thick in this specific application), thus saving printed material and also reducing part mass.

Nylon manifolds that are directly fabricated using LS must be printed thicker to have comparable strength.The vacuum bagging process cures the resin and hardens the carbon fiber while pulling excess resin out of the part (thus decreasing part weight). Additionally, the vacuum seals the pores present in the FDM-fabricated part as resin is pulled into the structure. Not only does this contribute to part strength, but it increases its resistivity to heat as the high-temperature resin permeates the part’s pores.Compared to lost core fabrication methods, this process requires fewer layers of carbon fiber as the ABS shell provides structure and strength. Additionally, the proposed process avoids many of the difficulties and frustrations typically encountered in the lost core method. The ABS shell ensures that the designed interior geometry is not compromised by poor composite adhesion and also precludes the need for the removal of a core.

The ABS shell ensures that the designed interior geometry is not compromised by poor composite adhesion and also precludes the need for the removal of a core. Furthermore, should gaps in the composite fabric occur during layup, there is no need to restart the process as the ABS shell provides sufficient strength for operation. Finally, the proposed process does not require the fabrication of a mold in which the pattern is created, reducing manufacturing cost and time.3 INTAKE MANIFOLD DESIGN AND MANUFACTURE3.1 Intake Manifold DesignAs the chosen manufacturing process removed prior geometric constraints, the student team was afforded the opportunity to redesign the intake manifold system geometry. In this redesign process, their efforts were directed towards two portions of the system: the length and geometry of the cylinder runners and the geometry of the plenum.The length and geometry of the cylinder runners were the first components designed.

Following the combined use of a dynamometer and computational fluid dynamics (specifically, CFdesign 10.0), atapered bent design of approximately 11 inches was selected. This length allows for additional air at the particular RPM range in which the car will operate, thus improving vehicle performance under these certain operating conditions. The team chose the bent design because of the need to fit the 11” runners within the packaging constraints of the vehicle.Runner geometry design decisions were dominated by the team’s focus on minimizing pressure losses at the transition from the plenum to the runner. The team discovered that recirculation could be eliminated by creating large inlets to the runners from the plenum and adding a taper to the runners (Figure 3). In addition, the tapered design increases air flow velocity into the cylinders, giving a better mixture of fuel and air to the engine and thus providing better combustion and overall performance.Figure 3.

Velocity vectors representing air flow as it travels from the plenum to the runners.In designing the plenum, it is important to achieve an even static pressure as this will cause the cylinders to pull the same vacuum, leading to even flow in each cylinder. In order to achieve this goal, a designer is typically faced with a tradeoff: even static pressures are easily achieved by larger plenum volumes, however this not only becomes difficult to package, it affects throttle response as a larger volume increases the amount of time for the system to reach an equilibrium pressure. With the ability to create a plenum of unlimited geometry, a design was developed that was tapered and could provide an even static pressure. This tapered design offsets the static pressure that is lost due to friction and other factors within the plenum since the taper increases the air velocity as it flows from the entry point to the end.

This decrease in size provides gives quick throttle response while keeping an even flow distribution to each cylinder, thus providing increased performance. The final intake manifold system geometry is shown in Figure 4.The improvement in the static pressure distribution over a similar aluminum design from the 2008 Virginia Tech Formula SAE team can be observed in Figure 5. The 2009 intake system design is similar to that of the 2008 team, featuring the same runner lengths, entry point, plenum volume, and basic layout. The 2008 system was manufactured from aluminum and has a slightly different geometry due to the limitations in using that manufacturing technique (specifically, the inability to taper the runners and the plenum). As seen in the graph, the static pressure drop across the plenum is a fraction of that of the 2008 design. It should also be noted that the static pressure is much lower in the plenum, which is evidence of a smaller pressure drop in the system to this point.

It should also be noted that the static pressure is much lower in the plenum, which is evidence of a smaller pressure drop in the system to this point. With an even static pressure, each cylinder pulls the same vacuum, ensuring that equal amounts of air are provided to each runner, thus increasing performance.Figure 4. Final intake manifold system model.1.21) i s p ( e r u s s e r P c i t a t S0.80.60.40.20123Position Based on Cylinder Number2008 Plenum 2009 PlenumFigure 5. Comparing the static pressure distribution of the 2008 and 2009 plenum designs.With the interior geometry of the intake system designed, the team turned their attention to the task of detailed design of the component. Specifically, features were added to the manifold model for sensor mounting and mounts for the fuel injector bosses.

Specifically, features were added to the manifold model for sensor mounting and mounts for the fuel injector bosses. The ability to integrate these mounts into the geometry reduced the overall manufacturing time, as it eliminated the time needed to manufacture aluminum pieces to be connected to the manifold.3.2 Intake Manifold ManufactureFollowing final design reviews and simulations, the CAD model was exported to the.stl format for fabrication on the lab’s FDM machine (Stratasys Dimension SST). Due to the constraints of the machine’s build volume (8”x8”x10”), the intake system geometry was modified such that the plenum elbow was built as a separate piece. The two pieces were built using the P400 ABS material. Soluble support material (P400-SR) was used to ensure the easy removal of support material from the hollow geometry of the manifold. Once fabricated, the parts were assembled together and bonded using an epoxy (Figure 6).4Figure 6.

Once fabricated, the parts were assembled together and bonded using an epoxy (Figure 6).4Figure 6. Final assembly of the intake manifold.With the assembly step complete, the student team proceeded with the application of the composite material onto the manifold. Three layers of carbon fiber fabric were applied to the manifold using a high temperature resin (PTM&W PT2520). This resin was chosen as it is able to survive the high temperatures (~120 F operating, up to 250 F during heat soak) that are endured from the operation of the engine. A vacuum bagging process was used to lay up the fabric onto the manifold. In addition to curing the resin and ensuring a proper fit of the fabric onto the manifold, the vacuum assisted in drawing the high temperature resin into the semi-porous ABS part.

In addition to curing the resin and ensuring a proper fit of the fabric onto the manifold, the vacuum assisted in drawing the high temperature resin into the semi-porous ABS part. In addition to adding some strength, the drawn-in of resin increased the ABS part’s resistivity to heat.The final manifold, featuring the completed composite layup and assembly of all mounts and sensors,is presented in Figure 7. The manifold, as mounted onto the Formula SAE car, is shown in Figure 8.Figure 7. Completed intake system, following the composite layup process and final assembly of sensors and mounts.The student design time was able to fabricate an end-use intake manifold in just under a week. The plenum and elbow were fabricated in 60 hours using the FDM machine (including support removal in a cleaning solution). An additional 3 days were required for surface and composite work.

The plenum and elbow were fabricated in 60 hours using the FDM machine (including support removal in a cleaning solution). An additional 3 days were required for surface and composite work. This manufacturing time is much less than the approximately 2 to 3 weeks that it would take to manufacture similar pieces in an aluminum, as they would require additional time for design, jig creation, and welding.Figure 8. The intake system as mounted onto the 2009 Virginia Tech Formula SAE car4. RESULTSAs stated in Section 1.2, the primary goal in the design of an intake manifold system is to improve the engine’s ability to efficiently and effectively produce torque and power through the even distribution of the combustion mixture evenly to each intake port. In order to estimate the flow distribution of the system, as well as visualize and analyze the flow to find areas of recirculation and separation that increase pressure drop, simulations were done using CFD.

These simulations were then compared with exhaust gas temperature (EGT) data. EGT data can be used to estimate the flow distribution; with more air the EGT data will be high, with more fuel the EGT will be low. The result from these simulations and tests can be seen in Figure 9 and Figure 10. The intake system designs from the 2009 and 2008 teams are compared in Figure 10.In Figure 10 it is evident that the design changes described in Section 3.1 have improved the flow distribution across the cylinders. These part-level improvements have transferred to improvements at the overall engine-level. Specifically, due to the new design, the engine package for 2009 car was able to achieve more torque than that of 2008. The differences in the torque curves of 2008 to 2009 as well as that of the most powerful natural aspirated (NA) engine at the 2007 competition can be seen in Figure 11.

The differences in the torque curves of 2008 to 2009 as well as that of the most powerful natural aspirated (NA) engine at the 2007 competition can be seen in Figure 11. Even though the engine package for 2009 had a lower peak torque, the torque curve is much more consistent over a longer range of RPMs, improving the drivability of the vehicle - a major design goal for the 2009 student team. The 2009 turbocharged engine also produced more torque throughout the complete RPM range over the naturally aspirated engine. The 2009 engine maintained a higher torque than the 2007 NA engine for a range of 6000 RPM compared to a 4700 RPM range for the 2008 engine.

The 2009 engine maintained a higher torque than the 2007 NA engine for a range of 6000 RPM compared to a 4700 RPM range for the 2008 engine. These improvements over previous designs are due largely to the new manufacturing process of the intake system, which enabled the design of a manifold that provided a more even flow distribution to each cylinder and a reduction of pressure losses.30.00)%(w o l F f o n o i t u b i r t s i D25.0020.0015.0010.005.000.0012 3 Cylinder Number4EGT DistributionCFD PredictionsFigure 9. Flow distribution into each cylinder based on CFD and EGT data.30.00)%(w o l F f o n o i t u b i r t s i D25.0020.0015.0010.005.000.0013 Cylinder Number 2009 Intake242008 IntakeFigure 10. Comparing the flow distribution of the 2009 and 2008 intake systems.Figure 11.

Comparing the flow distribution of the 2009 and 2008 intake systems.Figure 11. Comparing the torque of the 2009 turbocharged engine, the 2008 turbocharged engine, and the most powerful NA engine at the 2007 competition.5. CLOSUREIn this paper, the authors present a manufacturing process used to create an intake manifold for a Formula SAE car. Specifically, an intake manifold shell is first printed using Fused Deposition Modeling; composite material is then applied to the shell with a high temperature resin and a vacuum bagging process. The use of FDM provided geometric flexibility in the design of the manifold, while the use of composite material and high-temperature resin ensured that the system would have sufficient strength and heat-resistivity to survive the operating environment of the turbo-charged engine.The use of FDM provided the student design team the freedom to create a unique intake geometry that featured a tapered plenum and tapered runners.

These geometric features provided an even static pressure throughout the system, an equal charge to each cylinder, reduced pressure loss, and an increased air velocity into the cylinders - key goals in manifold design. The use of the additive manufacturing technology ensured the creation of a system free from sudden geometry changes, which reduce flow separation and increase pressure loss.The manifold that resulted from the process greatly improved the performance of the Formula SAE engine, as compared to a similar aluminum design. The new design provided an increase in torque over an extended RPM range as the manufacturing process provided the ability to create a complex geometry that provided equal air to each cylinder and reduced total pressure drop. Additionally, relative to the previous design, the total weight of the system was reduced from 2.875 lbm to 2.24 lbm (a 22% decrease).6. REFERENCES1.

Additionally, relative to the previous design, the total weight of the system was reduced from 2.875 lbm to 2.24 lbm (a 22% decrease).6. REFERENCES1. Jackson, S., 2008, “From Rapid Prototyping to Mass Customization,” Industrial Laser Solutions for Manufacturing, http://www.industrial-lasers.com/articles/article_display.html?id=333159, accessed July 2009.2. Gebhardt, A., 2003, Rapid Prototyping, Hanser.3. Abbott, P., 1998, “Rapid Prototyping,” Assembly Automation, Vol. 18, Iss. 1, Mini-feature.4. Dore, S., P. Lavallee, 1997, “Design and Fabrication of Intake Manifold for Formula SAE (Society of Automotive Engineers) Race Car,” Rapid Product Development Technologies, P. Boulanger (ed.), pp. 187-197.5. “Hi-Tech Hot Rod is the Ultimate Learning Lab,” 2005, ScienceDaily,http://www.sciencedaily.com/releases/2005/01/050125085013.htm., accessed July 2009.6.

sign and Manufacture of an Attachment Fitting for Transtibial Prosthetic Sockets Using Selective Laser SinteringDanny Burhan and Richard Crawford Laboratory for Freeform Fabrication The University of Texas at AustinAbstractThe focus of this work is using selective laser sintering to manufacture transtibial prosthetics sockets with compliant features to relieve contact pressure in sensitive areas. Each of these sockets requires an integrated attachment fitting to connect to the pylon and foot using standard hardware. Several design concepts of an attachment fitting are presented and compared. The design concepts were tested using a tensile test machine and analyzed using ground reaction force data to ensure a structurally sound connection.

The design concepts were tested using a tensile test machine and analyzed using ground reaction force data to ensure a structurally sound connection. The resulting design employs standard hardware while maintaining the integrity of the connection for a normal gait cycle.IntroductionIn United States there is a growing need to fabricate prosthetic sockets faster and moreeconomically, with more than 500,000 current lower extremity amputees and at least 60,000 new ones each year [1]. Conventional methods of fabricating prosthetic sockets are time consuming and expensive, and highly dependent on the skills of prosthetists, of whom there are fewer compared with the number of amputees.

The University of Texas at Austin, in collaboration with the Department of Rehabilitation Medicine at the University of Texas Health Science Center at San Antonio (UTHSCSA), is developing an alternative method of fabricating prosthetic sockets that can improve efficiency in prosthetic care, enhance comfort and fit, and reduce time and cost.The sockets are produced using Selective Laser Sintering (SLS), which creates complexobjects by adding and bonding materials in layers. SLS has the capability of forming any complex geometry without the need for elaborate machine setup or final assembly. Using SLS, prosthetic sockets can be created directly from digital shape information, eliminating the need for extensive molds, hand lamination and finishing procedures.

Using SLS, prosthetic sockets can be created directly from digital shape information, eliminating the need for extensive molds, hand lamination and finishing procedures. SLS also simplifies the construction of complex objects into a manageable, straightforward, and relatively fast process.Previous research on fabricating subject specific prosthetics sockets using selective laser sintering has been presented by Stevens [2] and Faustini [3]. The process was aimed specifically at the production of patellar- tendon bearing (PTB) sockets, modified with compliant features to relieve contact pressure at certain areas. Figure 1 shows a diagram of a prosthetic limb. The prosthetic limb consists of the socket with integrated attachment fitting, pylon and prosthetic foot. Designing an integrated attachment fitting to mate with standard hardware is the objective of this research.794ObjectivesFigure 1 shows that the attachment fitting provides the connection between the socketand pylon.

In previous work, the design of the attachment fitting used custom parts, which makes it harder to commercialize the product due to the increase in production cost. The purpose of this research is redesigning the attachment fitting to incorporate standard hardware while maintaining the integrity of the socket and a strong connection to the pylon. Utilizing the advantage of SLS to create complex objects without additional tooling, an integrated attachment fitting will be created.The research began by setting the design requirements, and then several designalternatives were developed and evaluated. The selected design concept was tested using a tensile test machine and analyzed using applied ground reaction forces (GRF) to ensure a strong connection is achieved.Figure 1: Prosthetic Limb.UTHSCSA Attachment Fitting DesignUTHSCSA has developed an alternative method of attaching the pylon to the socket.

Inorder to eliminate the use of custom parts, UTHSCSA uses a standard four hole European adaptor, as shown in Figure 2a. This type of adaptor has been used regularly to attach the pylon to conventional sockets and is easily obtainable from any prosthetic hardware vendor. The adaptor is then connected to a standard cylindrical connector (Figure 2b) that enables adjustment of the alignment angle. The adjustment is made by tightening and loosing four setscrews on the outer wall of the connector. Figure 2c is a representation of how the adaptor and connector are joined once the adjustment is made on the definitive socket.795Figure 2: (a) Standard Four Hole European Style Adaptor; (b) Standard Connector; (c) Adaptor Joined to Connector.For SLS sockets, UTHSCSA designed an integrated attachment fitting to mate with theadapter.

The design is simple; a square block is created on the bottom of the socket and united to the socket by using a “blend” operation in the CAD software, creating one continuous smooth surface. In order to secure the adaptor, two small aluminum blocks with internal threads are inserted and tightened using standard 6mm bolts (Figure 3). These aluminum inserts are used because UTHSCSA did not believe that creating direct tapped holes in DuraformTM PA would be possible. Although the UTHSCSA attachment fitting design utilizes standard prosthetic hardware, the design is not completely standard. Even though the cost of manufacturing the custom aluminum blocks is inexpensive, it is not widely available off the shelf.

Even though the cost of manufacturing the custom aluminum blocks is inexpensive, it is not widely available off the shelf. Moreover, the UTHSCSA attachment fitting has not been analyzed or optimized.Figure 3: (a) UTHSCSA Attachment Fitting; (b) Aluminum Block Inserts.Attachment Fitting Design RequirementsThe requirements for design of the attachment fitting were set by Gordon Bosker and BillRogers [4] from UTHSCSA. As mentioned previously, the usage of all standard hardware is very important for the success of the design. Using standard hardware will improve marketability of the SLS socket throughout the United States and the world. The following are design requirements for the attachment fitting:7961. The design should use the standard four hole European style adaptor (Figure 2a) for the design, since this adaptor will provide adjustment for the alignment angle.2.

The design should use the standard four hole European style adaptor (Figure 2a) for the design, since this adaptor will provide adjustment for the alignment angle.2. The inner profile surface of the socket should be maintained, meaning the fitting should not cause any alterations. The inner profile has been rectified by the prosthetist so that contact pressure is distributed in areas that are less sensitive.3. The overall shape of the socket should be aesthetically appealing. The attachment fitting should blend smoothly to the socket and should not be too bulky, as this will increase the overall weight of the socket.4. The attachment fitting should allow the adaptor to be removed when necessary, which means that the adaptor should not be glued permanently to the socket.5. The use of other standard prosthetic hardware is recommended. This hardware usually is attached to the conventional socket using epoxy.

The use of other standard prosthetic hardware is recommended. This hardware usually is attached to the conventional socket using epoxy. Other standard or universal hardware that is easily accessible is also recommended.6. The attachment fitting should withstand a normal walking force.7. The design should not require manufacture of non-standard parts.8. Fabrication of the attachment fitting should be compatible with SLS.9. DuraformTM PA should be used as the material for the attachment fitting.Attachment Fitting Concept GenerationBased on the design requirements, several attachment fitting concepts were created usingstandard hardware. Figure 4 shows several of these possibilities. Figure 4a shows two of the aluminum inserts that are currently used by UTHSCSA. Figure 4b is a 6mm flange nut, a standard metric nut for a 6mm prosthetic bolt. Figure 4c is a T-insert, usually used in wood, but also a possible insert for DuraformTM PA.

Figure 4b is a 6mm flange nut, a standard metric nut for a 6mm prosthetic bolt. Figure 4c is a T-insert, usually used in wood, but also a possible insert for DuraformTM PA. Figure 4d is a helicoil, normally used for thread repair. The helicoil works by expansion of its threads after a 6mm bolt is inserted. Figure 4e shows a tap that would be used to create a threaded hole directly in the SLS material. No previous studies have been performed to test the strength of direct tapped holes in DuraformTM PA.797Figure 4: Standard Metal Inserts.Each of the concepts using standard hardware was refined to determine exactly how thehardware would mate with the socket and the adaptor. As a start, the base of the UTHSCSA attachment fitting (measuring 60mm x 60mm x 16mm) was used (Figure 3a), since it was known to provide solid support for the adaptor. For the flange nut, the attachment fitting is designed around the shape of the nut itself.

For the flange nut, the attachment fitting is designed around the shape of the nut itself. Figure 5b shows the proposed attachment fitting using the flange nut design. The flange nut is inserted from the side. Once the nut is inside, the wall around the flange nut secures the nut as the bolt is tightened. The adaptor is then installed the same way as it is in UTHSCSA design, directly underneath the socket.Just like the flange nut design, the T-insert attachment fitting is designed around theinsert itself (Figure 5c). The T-insert is inserted from the bottom and glued using epoxy. The spikes in the t-insert will prevent any rotational movement when the bolt is tightened. A feasibility study of gluing metal to DuraformTM PA was performed to ensure that a bond to DuraformTM PA can be formed. The study consisted of gluing the flat side of the t-insert to DuraformTM PA using an all-purpose epoxy. Before gluing, the surface of the plate was cleaned thoroughly to remove loose powder.

The study consisted of gluing the flat side of the t-insert to DuraformTM PA using an all-purpose epoxy. Before gluing, the surface of the plate was cleaned thoroughly to remove loose powder. The result shows that gluing metal objects to DuraformTM PA is a feasible alternative.Attachment fitting designs for the helicoil insert and the direct tapped holes are straightforward (Figure 5d). The base of the socket where the adaptor is connected is the same as for the flange nut and T-insert designs. Both of the designs require drilling holes for the 6mm helicoil insert and the 6mm direct tapped hole. A special tap is required for inserting the 6mm helicoil and the required hole for this tap is 0.25”. Once the holes have been tapped, the helicoil is then inserted using a special tool. After the helicoil reaches the desired depth, the special tool cuts off the tip of the helicoil, clearing the path for the bolts to be inserted.

After the helicoil reaches the desired depth, the special tool cuts off the tip of the helicoil, clearing the path for the bolts to be inserted. For the direct tapped hole, a regular 6mm tapped hole is required. Once the hole has been tapped the bolts are inserted normally.Four of these refined designs were presented to UTHSCSA staff. All of the designs are simpler and use standard commercial hardware. UTHSCSA staff recommended testing each of the concepts. Each of these concepts was compared to the UTHSCSA design. The results of these tests are presented below.798Figure 5: (a) UTHSCSA Attachment Fitting Design; (b) Flange Nut Attachment Fitting Design; (c) T-insert Attachment Fitting Design; (d) Helicoil and Direct Tapped Attachment Fitting Design.Concept SelectionTo test and compare the concepts, a MTS 810 tensile test machine was used. The machine has two spring-loaded grips that are normally used to hold “dogbane” tensile specimens.

The machine has two spring-loaded grips that are normally used to hold “dogbane” tensile specimens. The grip has a maximum opening of 17mm; therefore, the specimen needs to have a maximum width of 17mm. To accommodate the width of the grip, a custom specimen for each concept was needed (Figure 6). Each specimen was created based on the location of probable failure. Analyzing the UTHSCSA and flange nut design, the strength of the design depends on the sidewall of the holes. For the T-insert, the strength depends on the bonding strength of the epoxy glue, and for the helicoil, the strength depends on the wall thread of the helicoil. Therefore, testing one corner of a complete attachment fitting, which includes just one hole, is sufficient since this reflects the strength of each individual design.For the test, a total of twenty-five specimens were built, five specimens for each concept. The testing follows ASTM Standard D638-02a procedure for testing polymer material in tension.

The testing follows ASTM Standard D638-02a procedure for testing polymer material in tension. Although the specimen dimensions do not comply with the specifications of the standard, the procedures in the testing do follow steps in the standard. A feed rate of 5mm/min was chosen based on the standard. The results of the testing are summarized in Table 1. The table lists the tensile force required to break the specimens in pounds-force (lbf). It also lists the average and the range for each concept. The range for the design is included to indicate the consistency of the test. As can be seen in the table, two of the direct tapped specimens did not provide good799results and were not recorded. This was probably due to poor tapping, which indicated that the same mistakes could occur if the direct tap design was chosen as the selected concept.

This was probably due to poor tapping, which indicated that the same mistakes could occur if the direct tap design was chosen as the selected concept. For this reason, the direct tap design was immediately eliminated as one of the concepts.Table 1: Tensile Test Force Results (lbf).Test 1 2 3 4 5 Average RangeFlange UTHSCSA T-Insert 566.81 1401.91 1304.42 756.50 1370.93 1166.12 659.01 1433.65 1230.36 466.30 1433.65 1237.92 157.20 1445.75 1203.91 521.16 1417.18 1228.54 599.31 74.82 138.30Helicoil 1154.78 1123.04 1287.80 1164.61 1165.36 1179.12 164.75Direct Tapped 365.78 442.11 487.46 ― ― 431.78 76.33Figure 6 shows the failure location for each specimen. As predicted, all UTHSCSA and flange specimens failed at the sidewall, since this is the weakest point of the design.

As predicted, all UTHSCSA and flange specimens failed at the sidewall, since this is the weakest point of the design. For the T- insert design, the failure occurred when the insert was fully extracted out of the base, whereas the helicoil and direct tapped failed when the DuraformTM PA threaded wall was completely stripped out. The result of the testing indicates that the UTHSCSA design is the strongest. Also, the range of the UTHSCSA concept is smallest, indicating the test was consistent. Theoretically, the flange design should be stronger than UTHSCSA design since it has a larger sidewall cross- sectional area, but the results from the test indicate the contrary. This issue is discussed below. The results for the T-insert design are very inconsistent since these data have the largest range. The strength of the T-insert design is highly dependent on the amount of epoxy in contact with the wall of the hole and the insert.

The strength of the T-insert design is highly dependent on the amount of epoxy in contact with the wall of the hole and the insert. The amount of epoxy that was applied was very difficult to control in every specimen, and resulted in inconsistent test results. The test results for the helicoil design were consistent and showed good strength. Even though the helicoil is not the strongest design, it does prove that the helicoil improves the strength of a threaded hole in DuraformTM PA compared to a direct tapped hole.Figure 6: Condition of Test Specimens after Testing.800The test results do not give a definitive answer in terms of which concept is strongest.The UTHSCSA and flange designs were stronger than any other designs, but because of the counterintuitive results from the test, further testing was needed to determine what actually happened. The flange nut design was suspected to have lower strength because of its build location in the SLS workstation.

The flange nut design was suspected to have lower strength because of its build location in the SLS workstation. All of the flange nut specimens were built in the front part of the build cylinder, whereas the UTHSCSA specimens were built towards the back. Build location can be a factor in determining strength, since there are several heaters in the SLS workstation and the temperature of these heaters can vary, thus affecting the strength of the part [5]. To test this hypothesis, a new set of UTHSCSA and flange specimens were fabricated with the build locations switched.Because of problems with the MTS 810 tensile machine, an Instron Tensile Machine wasused for the second set of tests. Since the Instron’s grip is smaller and narrower than the MTS 810 grip, the previous specimens did not fit in the Instron’s grip. To verify whether build location plays a role, a simpler approach was pursued. Instead of testing the concept specimens made earlier, standard dogbanes were tested.

To verify whether build location plays a role, a simpler approach was pursued. Instead of testing the concept specimens made earlier, standard dogbanes were tested. Ten dogbanes were built and tested based on ASTM Standard D638-02a. Five of them were built in the front of the build cylinder and the other five in the back. The results of the test are summarized in Table 2.Table 2. Dog bones Test Results.Specimen Front - 1 Front - 2 Front - 3 Front - 4 Front - 5 Average RangeMax Load (lbf) Specimen Back - 1 Back - 2 Back - 3 Back - 4 Back - 5 Average Range175.35 182.10 211.32 218.06 220.31 201.43 47.96Max Load (lbf) 254.03 240.55 224.81 200.08 220.31 227.96 53.95The test results show that the specimens built in the back of the build cylinder were stronger than the ones in the front. These results support the explanation of why the flange design specimens were weaker. The flange design should perform as well as the UTHSCSA or even better.

These results support the explanation of why the flange design specimens were weaker. The flange design should perform as well as the UTHSCSA or even better. Based on the result of the dog bone tensile test results, the flange nut design was selected as the attachment fitting for the SLS prosthetic socket. Under ideal conditions, the flange nut design should perform slightly better than UTHSCSA design.Testing the Complete Attachment FittingAfter selecting the flange nut design as the attachment fitting, the strength of the wholeattachment fitting was evaluated, assuming the attachment fitting is the weakest point in the prosthesis. Figure 5b was used as a starting design for testing the attachment fitting. Since other areas of the socket that are more susceptible to failure than the fitting, testing the whole socket with the attachment fitting would yield information about the strength of the fitting design. Therefore a specific test specimen was needed.

Therefore a specific test specimen was needed. The complete attachment specimen needs to fit in the available tensile test machine, the MTS 810 machine. For this test, using the grips was not801possible since the width of the attachment fitting itself is greater than 17mm. Figure 7a shows the complete attachment fitting specimen and Figure 7b shows the setup for the test. The specimen was made by taking the base feature in Figure 5b and blending it with flange nut design. The upper portion of the specimen accommodates a one-inch threaded shaft coming out of the test machine (See Figure 7b). This shaft is then secured using a one-inch nut and washer. A larger surface area was created for the upper feature to accommodate the one-inch nut and to ensure that the upper feature did not fail first.

A larger surface area was created for the upper feature to accommodate the one-inch nut and to ensure that the upper feature did not fail first. The bottom part of the specimen was connected to a custom plate with a threaded shaft secured to the specimen using four 6mm bolts.The test was performed using a 5mm/min feed rate, and the force to break the specimenwas 5799.61 lbf (25,797.95 N). This force was more than four times the average individual specimen tested previously, which was 4914.16 lbf. This result is expected since the cross- section of the full attachment fitting is larger than the area of the individual specimen. The test also indicates the location of failure is on the sidewall where the flange nuts are located. Figure 7c also shows the condition of the attachment fitting after failure. The fracture is conical and located at the holes, indicating that stress concentrations appeared before the specimen finally failed.

The fracture is conical and located at the holes, indicating that stress concentrations appeared before the specimen finally failed. The result of this test was then compared with ground reaction force data of an amputee wearing a traditional prosthetic socket. Using the GRF data, the attachment fitting was evaluated to determine whether it is strong enough to withstand normal walking forces.Figure 7: (a) Complete Attachment Fitting Specimen; (b) Test Setup; (c) Specimen Condition after Testing.Analysis of Ground Reaction ForcesThe ground reaction forces are forces between the foot and the ground as a person walks.These forces are vectors with components in the xyz directions that are measured using an instrumented force plate. A patient wearing a prosthetic socket walks on the runway and sensors inside the force plates record the forces as the foot is in contact with the plate. The ground802reaction forces are recorded using a global coordinate system fixed to the ground.

The ground802reaction forces are recorded using a global coordinate system fixed to the ground. Vertical forces are represented by Fz, anterior-posterior forces by Fx, and transverse (side to side) forces by Fy (see Figure 8a for representation of the forces). These forces are recorded at the foot base, and to be useful for this research, the GRF data needs to be resolved into individual components. For the purpose of designing the attachment fitting, the most critical forces are forces that are perpendicular to the foot, since these forces create a bending moment on the socket and can cause the socket to fail. The shear forces are carried by the bolts and are not as likely to cause failure. The only GRF force that is always perpendicular to the shank is the transverse force, Fy.

The shear forces are carried by the bolts and are not as likely to cause failure. The only GRF force that is always perpendicular to the shank is the transverse force, Fy. The vertical (Fz) and anterior-posterior (Fx) forces have both perpendicular and parallel force components (Figure 8a shows the perpendicular force components) and therefore needs to be further calculated. In order to extract the perpendicular forces, the walking angle is needed. The walking angle is defined as the angle between the shank and the ground (Figure 8b).

In order to extract the perpendicular forces, the walking angle is needed. The walking angle is defined as the angle between the shank and the ground (Figure 8b). Since the GRF data does not contain the angle information, an alternative way of incorporating the walking angle is needed.a.Fzb.αHipα β γ= Hip Angle = Knee Angle = Walking Angle + angle = counterclockwise - angle = clockwiseFx perpendicular FxFy = Fy perpendicularFz perpendicularShankβγFigure 8: (a) GRF Perpendicular Forces Component; (b) Illustration of Interpreting Waking Angle.The method for obtaining the walking angle data is based on [6], which presents hip, knee, and ankle angles for a normal person walking. Since the interest of this study is for an amputee wearing a prosthesis, the ankle angle is not needed. Figure 8b shows how the walking angle is acquired.

Since the interest of this study is for an amputee wearing a prosthesis, the ankle angle is not needed. Figure 8b shows how the walking angle is acquired. The hip angle (α) is the angle between the hip and vertical body movement, while the knee angle (β) is the angle between the extension of the hip and the shank. Therefore, the walking angle (γ) is the angle between the ground and the shank, and is computed as a simple function of α and β. In [6], the angle is presented over a complete gait cycle, which includes the stance and swing phases. The stance phase, which is about 65% [7] of a complete cycle, is the only phase that is needed for this analysis, so only 65% of the angle data was used.In order to fully extract the perpendicular forces, the correlation between the GRF dataand the walking angle was needed. Four sets of GRF data from amputees wearing prosthetic sockets were acquired from the UTHSCSA Gait Lab.

Four sets of GRF data from amputees wearing prosthetic sockets were acquired from the UTHSCSA Gait Lab. Each person’s GRF data contained 32 data points that represented forces in the xyz directions. For each data point, a walking angle was needed to extract the corresponding perpendicular force. Therefore, the walking angle was calculated over 32 angles.803Once walking angles were acquired, perpendicular components of vertical and anterior-posterior forces were calculated. With all the perpendicular forces known, a resultant force was then obtained by calculating the resultant of the perpendicular vertical force and anterior- posterior force. This resultant force was combined with the transverse force. The total resultant of these three forces was used to evaluate the attachment fitting under bending.

This resultant force was combined with the transverse force. The total resultant of these three forces was used to evaluate the attachment fitting under bending. A summary of the maximum for each force and the total perpendicular force is presented in Table 3.Table 3: GRF Maximum Force.Fx (N) Fy (N) Fz (N) Total Perpendicular Resultant Force (N) 87.6 Amputee 1 79.7 Amputee 2 Amputee 3 78.2 Amputee 4 169.4134.7 107.8 145.0 200.6910.7 638.5 974.4 1001.9353.4 257.0 443.2 422.3Relationship of GRF to Tensile Test ResultsAfter extracting the maximum resultant forces from the GRF, the next step wascorrelating the forces to the test results of a complete attachment fitting. Figure 9a shows a diagram of the attachment fitting. The diagram is arranged to represent an attachment fitting, a connected adaptor, and a pylon.

Figure 9a shows a diagram of the attachment fitting. The diagram is arranged to represent an attachment fitting, a connected adaptor, and a pylon. The results of the analysis predict whether the attachment fitting is strong enough under normal walking forces.As indicated in Figure 9c, a bending force, which is the maximum resultant force from the GRF analysis, is applied to the end of the pylon and, as the force is applied, it will pivot at point A. Ftop and Fbottom represent forces on the adaptor as the bending moment force is applied. Therefore the reaction forces from the adaptor are the tension forces that the attachment fitting experiences. The perpendicular distances between point A to the forces are represented by a, b, and c.

The perpendicular distances between point A to the forces are represented by a, b, and c. Solving for the values of Ftop and Fbottom gives the total tension force applied to the attachment fitting.a.b.c.A T T A C H M E N TF I T T I N GA D A P T O RBending ForcePYLONb2 Vtop2 VbottomcFAA D A P T O R2 FtopBending ForcePYLON2 Fbottom aAδθAFigure 9: (a) Attachment Fitting Diagram; (b) Free Body Diagram of Attachment Fitting; (c) Rotation Angle.804From the free body diagram in Figure 9, with the adaptor tending to rotate about point A,the strains on the top and bottom bolts areδTop= b×δθδBottom= c×δθ,where δθ is the rotation angle between the bottom of the attachment fitting and the top of the adaptor.

The force on each bolt is a function of the strain:FBolts = Ff f δ=(strain ( Top ) ( BottomtopFf δ=bottom))Since δθ is the same for both bolts and the strains are proportional to the distances to the pivot point A, the forces can be assumed proportional:Ftopb ×= cFbottomSumming the moments at point A givesFbending∑ = 0 AM a =× ×2(Ftop×b)+2(×Fbottom×c)Solving equations (1) and (2) simultaneously for the given bending load gives the values for Ftop and Fbottom. The total tensile force is then 2Ftop + 2Fbottom. The total tensile force of the attachment fitting was computed using the measured distances and maximum resultant force from GRF analysis.

The total tensile force is then 2Ftop + 2Fbottom. The total tensile force of the attachment fitting was computed using the measured distances and maximum resultant force from GRF analysis. Table 4 summarizes the maximum tensile force the attachment fitting experiences.Table 4: Total Tension Force of the Attachment Fitting due to Bending Force.Ftop BoltFbottom BoltFtotalAmputee 1 Amputee 2 Amputee 3 Amputee 4Newton 1440 953 1806 1567lbf Newton 324 214 406 352255 169 320 278lbf Newton 57 38 72 623390 2245 4252 3689lbf 762 505 956 829The results from the analysis show that the highest maximum tensile force applied to theattachment fitting is only 956 lbf. This force is far below the force needed to break the attachment fitting. Even though all four bolts experienced the same maximum force (in this case805(1)(2)4 times 406 lbf), which is 1624 lbf, the attachment fitting is still strong enough.

Even though all four bolts experienced the same maximum force (in this case805(1)(2)4 times 406 lbf), which is 1624 lbf, the attachment fitting is still strong enough. Using the value of 1624 lbf as an upper limit for tensile force and incorporating a safety factor of 3 yields a tensile force of 4872 lbf, which is significantly smaller that the force needed to break the attachment fitting (5799.61 lbf) in the experiment. This indicates the attachment fitting will be strong enough under normal walking loads.From the free body diagram in Figure 9b, the shear force on each of the four bolts is the same. By summing the forces in the y direction, the shear force for each bolt can be calculated as followsVbolt=∑ = 0 yF BendingFor ce NVbolt = BendingFor 4×42.443 4=ce=8.110N,(3)where the bending force is equal to the calculated perpendicular force in Table 4.

Using the calculated values of the largest maximum tensile force in Table 4.2 and the shear force for each bolt, the equivalent stress of the bolt can be calculated using the distortion energy equation [8]:2 3τσσ2=e+,(4)where the tensile stress,=σ1806 AN, the shear stress,=τN8.110 A, and A = 20.1 mm2 for a 6mmbolt [8]. Substituting these values into equation (4) yields the equivalent stress for the bolt, which is which means that using SAE class 4.8 bolts with a proof load of 310 MPa [8] will be sufficient to ensure that the bolts will not fail.3.90=eσ.MPaIncorporating a safety factor of 3 yields a tensile stress of 270.9 MPa,ConclusionThe primary result of this research is the design and testing of a new integrated attachment fitting for transtibial prosthetic sockets fabricated by selective laser sintering. The flange nut attachment fitting provides an economical, standard, and strong connection to the pylon.

The flange nut attachment fitting provides an economical, standard, and strong connection to the pylon. Flange nuts are easily obtainable and inexpensive compared to the UTHSCA design. The design should help marketability of SLS prosthetic sockets. The tensile test results and ground reaction force analysis indicate that the flange nut design is strong enough to withstand normal gait forces.References[1] Rogers, W. E., Crawford, R. H., Beaman, J. J., and Walsh, N. E., 1991, “Fabrication of Prosthetic Sockets by Selective Laser Sintering,” 1991 Solid Freeform Fabrication Symposium Proceedings, Marcus, H. L., Beaman, J. J., Barlow, J. W., Bourell, D. L., and Crawford, R. H., eds., Austin, TX, August 12-14, 1991, pp. 158-163.806[2] Stephens, S.

SIGN AND MANUFACTURING OF BONE ANALOG MODELS FOR THE MECHANICAL EVALUATION OF CUSTOM MEDICAL IMPLANTSTimothy J. Horn*, Ola L.A. Harrysson*, Jeffrey P. Little†, Harvey A. West II*, Denis J. Marcellin-Little†Edward P. Fitts Department of Industrial and Systems Engineering †Department of Clinical Sciences, College of Veterinary Medicine North Carolina State University, Raleigh NCAbstractThe performance of orthopedic implants is often evaluated using cadaveric bonespecimens. The high inter-specimen variability of cadaveric bone properties requires large sample sizes to obtain statistical significance. With recent focus on custom implants manufactured using direct metal freeform fabrication techniques, the need for a customized bone analog model is recognized. Data for bone geometry and internal structure were obtained from computed-tomography imaging.

Data for bone geometry and internal structure were obtained from computed-tomography imaging. Traditional rapid prototyping techniques are then used to generate the rapid tooling from which composite bones that mimic the properties of the real bone can be duplicated. This work focused on the manufacturing process of bone analog models.1.0 IntroductionAdditive manufacturing techniques that facilitate the direct fabrication of metal components such as electron beam melting (EBM) or selective laser sintering (SLS), are poised to revolutionize orthopedic implant and medical device manufacturing. Three-dimensional geometry data obtained with medical imaging systems such as computed tomography (CT) or magnetic resonance imaging (MRI) facilitate the design of custom implants and constructs that conform to a given individual's anatomy9.

These custom implants are directly fabricated using the EBM or SLS processes, and may incorporate features to optimize healing and long term success by precisely controlling mechanical properties, material properties, surface topography, and implant structure. Because custom implants can be made to conform to the individual patients anatomy, orthopedic procedures can be carried out with little or no modification to the underlying bone structure thereby optimizing the bone-implant interface11.

The additive nature of EBM and SLS also facilitates the generation of engineered lattice structures which can be optimized both for bone fixation, and for the transfer of mechanical loads through the implant and into the healthy bone tissue, thus minimizing stress shielding and bone resorption1,9.Despite the clear advantages of utilizing custom implants fabricated using direct additive manufacturing techniques, there is still a great deal of difficulty associated with the mechanical testing and evaluation of such implants. Traditionally, the pre-clinical biomechanical testing of orthopedic implants is carried out in vitro using materials testing equipment. Testing is typically comparative in nature, and cadaveric bones are often used as a substrate because they are available as matched pairs15.

Testing is typically comparative in nature, and cadaveric bones are often used as a substrate because they are available as matched pairs15. Furthermore, the size, shape, and properties of cadaveric bones typically allow an acceptable approximation of in vivo situations.However, the high cost and lack of availability of cadaveric bones coupled with biological degradation and the potential transmission of infectious diseases, complicates the use of cadaveric specimens for the mechanical evaluation of orthopedic implants 6,23. Although there is a general symmetry between right and left pairs of bones from the same cadaveric specimen, the high interspecimen variability between cadaveric specimens necessitates the use of large sample sizes to detect small differences between different implants or surgical methods21. In one study, the interspecimen variability of the mechanical properties of cadaveric specimens reached 100% of the mean2.

In one study, the interspecimen variability of the mechanical properties of cadaveric specimens reached 100% of the mean2. Further, paired comparisons using cadaveric specimens are limited to the evaluation of a single independent parameter between the two groups, therefore, the sample size problem is greatly exacerbated when it may be desirable to test the differences between more than two configurations of an implant.To overcome these disadvantages, materials such as wood18, plastic8, and polyurethane foam16 are sometimes used to model the mechanical properties of both cancellous and cortical bone. These models can reduce variability in testing (and therefore the required sample size) but lack realistic bone geometry. Bone models that replicate the shape of average human long bones are now commercially available, the latest iteration of which utilizes discontinuous short-fiber reinforced resin molded around a foam core13.

Although data regarding the formulation of the fiber reinforced composites are lacking in the literature, these models have demonstrated some mechanical properties that model natural human bones and at the same time exhibit comparatively low interspecimen variability2,13. The geometry of these bones is derived directly from a cadaveric specimen representing the average size and weight of a human male12.While a standardized analog bone model has many advantages, particularly when comparing data across multiple studies, this approach neglects the patient-specific geometry of custom implants fabricated with additive manufacturing technologies, making the meaningful evaluation of such implants impossible. In order to evaluate the in vitro performance of custom implants, patient-specific bone analog models are required that are both geometrically and mechanically accurate.

In order to evaluate the in vitro performance of custom implants, patient-specific bone analog models are required that are both geometrically and mechanically accurate. This paper addresses these issues and presents; (1) the development of materials that can model the intensive material properties of cortical and cancellous bone, (2) the development and manufacture of patient-specific bone analog models of a canine radius utilizing CT data, rapid prototyping, and a rapid tooling, and (3) an examination of the potential for the direct fabrication of custom bone analog models using rapid prototyping techniques.2.0 Development of Bone Analog MaterialsBone provides basic structural support for the body. Its internal architecture is organized in hierarchical fashion from constituent molecules to distinctive tissue types.

Its internal architecture is organized in hierarchical fashion from constituent molecules to distinctive tissue types. At the lowest level bone can be thought of as a composite material made up of three primary components; the fibrous protein type I collagen, which acts as an organic matrix; a mineral phase called carbonated hydroxyapatite, and water24. At the next level these components can be organized into groups of osteons that, combined, form the bulk of the dense outer shell of the bone called cortical bone, or into the less dense packets of lamellae that form cancellous bone3. The combination and organization of these components give rise to the unique material properties exhibited by bone.On the highest, macroscopic level, the long bones of the appendicular skeleton consist ofa long hollow central shaft called the diaphysis, with two wider portions on each end called the epiphyses. The illustration in Figure 1 shows a simplified schematic cross section of a canine radius.

The illustration in Figure 1 shows a simplified schematic cross section of a canine radius. The primary component of the diaphyseal portion of the bone is cortical bone, while the epiphyses contain a higher portion of cancellous bone surrounded by a thin cortical shell. Bone is also graduated, with compositions and configurations that differ from one geometric region to another; in particular, the transition from the cortical outer shell to the spongy, inner, cancellous bone is of particular interest to this study19.Figure 1: Schematic cross section illustrating some of the macroscopic components of a canine radius.Although there is a large body of literature dedicated to the study of the mechanicalproperties of the low level constituents of bone, such as osteons or lamellae, most research on bone material properties has focused on cortical and cancellous bone17.

Therefore in the development of analog materials to model bone, this study focuses on these two components.2.1 Cortical Bone Analog DevelopmentThe reported values for the mechanical properties of cortical bone in the literature varysignificantly; this is in part due to the natural variability of the material, as well as the variability testing methods from one study to another. In some studies whole bones were tested, typically in three or four point bending19. Other methods involve resting small sections of cortical bone were tested in tension, compression or indentation. In general the reported values for the modulus of elasticity of cortical bone for several mammalian species range from about 7 GPa to 34 GPa (1015 ksi to 4931 ksi )During the selection of a material to model the cortical bone, similarity to real corticalbone, availability, ease of preparation, ease of manufacturing, and the ability to vary the properties to match patient-specific data were considered.

The mechanical properties of a host bone have been successfully estimated from quantitative computed tomography scans5,20. Although it was not investigated in this study it may also be useful to match the properties of unhealthy or osteoporotic bone. For these reasons a fiberglass reinforced thermosetting epoxy resin was chosen for this study.This material is easily mixed and injection molded, and the properties can be varied by adjusting the fiber content, matrix material, fiber material, fiber length, etc. In addition, epoxy- fiberglass mixtures have been used to replicate human cortical bone in several studies12,13,23. No data are present in the literature regarding the formulation of these materials.

In addition, epoxy- fiberglass mixtures have been used to replicate human cortical bone in several studies12,13,23. No data are present in the literature regarding the formulation of these materials. A mathematical model was used to determine the possible range of mechanical properties attainable through this method (given the constraints on processing and manufacturing), as well as to predict the modulus of elasticity of the cortical bone analog material.Several mathematical models have been developed for the estimation of the tensilemodulus of elasticity for discontinuous fiber composites, but the Cox model is widely used and regarded to be the simplest26. This model averages the elastic constants of the material over all possible fiber orientations4.

This model averages the elastic constants of the material over all possible fiber orientations4. This is represented in Equation 1 as a modified version of the rule of mixtures.Equation 1:Ec=φχχ 2 f1Ef+φE mmWhere, Ec, Ef, and Em are the tensile elastic moduli of the composite, the fiber and thematrix respectively. φf and φm are the fiber volume fraction and the matrix volume fraction. The variable χ1 represents the fiber orientation efficiency factor which is equal to 1 for axially aligned discontinuous fibers and 1/5 for randomly oriented fibers12. Details for the determination of the fiber orientation efficiency factor are discussed in several sources4,7.

Details for the determination of the fiber orientation efficiency factor are discussed in several sources4,7. The variable χ2 represents the fiber length correction factor as is given by Equation 2 – 4 below; as the fibers become longer this value approaches 1.Equation 2:−=χ12tanh( nana)Equation 3:n=1−Ef2 Gln(mπ 4 φ f);Gm=1(2E m + ν m)Where: Gm is the shear modulus of the matrix, and νm is the Poisson ratio of the matrix.Equation 4:a/=dLWhere a is the aspect ration of the fiber, L is the length of the fiber and d is the diameter of the fiber. As demonstrated by Gibson7, the fiber orientation factor has a much greater impact on the mechanical properties of the composite than does the fiber length.

As demonstrated by Gibson7, the fiber orientation factor has a much greater impact on the mechanical properties of the composite than does the fiber length. Therefore the case of axially aligned discontinuous fibers provides a reasonable upper bound for the modulus of the composite.The rule mixtures can also be applied to determine a reasonable lower bound for the composite modulus by considering the case of transversely aligned fibers (in which both the fiber and the matrix are under the same isostress condition) as given in Equation 5.Equation 5:Ect=EEE m φ + mff Eφ fmThis set of equations provides an upper and lower bound on the tensile modulus of adiscontinuous fiber composite. For the cortical bone analog, epoxy resin and hardener (Fiberglast Corporation, Series 2000 and 2060 respectively) mixed in a 100:23 ratio (by weight) was used as the matrix material. The average modulus of elasticity of the cured resin was 3.091 GPa (448.3 ksi).

The average modulus of elasticity of the cured resin was 3.091 GPa (448.3 ksi). The fibers used were 793µm (0.0313 inch) in length S-glass fibers with a diameter of 20µm (0.007 inch). The average modulus of elasticity of the glass fibers was 89 GPa (12908 ksi).

The fibers used were 793µm (0.0313 inch) in length S-glass fibers with a diameter of 20µm (0.007 inch). The average modulus of elasticity of the glass fibers was 89 GPa (12908 ksi). Inserting these values into the mathematical models, Figure 2 shows a plot of the predicted tensile modulus of the cortical bone analog versus fiber volume fraction for various conditions of fiber alignment.45.040.0Random ly Oriented Fibers Aligned Fibers Low er Bound35.0) a P G30.0(s u25.0lu d o M e l i s n e T20.015.010.05.00.000.10.20.30.40.5Volume Fraction of FiberFigure 2: Plot illustrating the relationship between the volume fraction of fiber and the predicted tensile modulus for various conditions of fiber alignment (using 20µm diameter, 793µm long fibers).In order to test the properties of the cortical bone analog material, tensile specimens wereprepared in accordance with ASTM standard D638-08.

A silicone rubber mold (Smooth-on Moldmax 30TM) was used to fabricate the samples to the required dimensions. The epoxy resin and hardener were mixed with the glass fibers under vacuum for 10 minutes and then injected into the silicone rubber mold. After injection, the mold was placed in a pressure chamber at 482.6 kPa (70 psi) for a period of one hour, until the epoxy cured. The samples were then post cured in a convection oven at 75ºC (167ºF).Twenty-eight tensile specimens were prepared with four different fiber loadingconditions; 0%, 33% , 50% , and 60% fraction by weight (0%, 18%, 30%, and 39% by volume). Tensile testing was carried out on each of the formulations using an ATS 1620C testing machine with square, serrated grips, (Applied Test Systems, Butler, Pa.).

Tensile testing was carried out on each of the formulations using an ATS 1620C testing machine with square, serrated grips, (Applied Test Systems, Butler, Pa.). An extensometer with a gauge length of 50.8 mm (2.0 inches) (Epsilon Technologies Corp, Jackson, WY) was affixed to the narrow section of the samples, and testing was carried out at a loading rate of 1.27 mm/min (0.05 in/min). The modulus of elasticity of each sample was calculated from the linear portion of the stress-strain plot.The results were compared within groups for all loading regimens using paired t-tests after confirmation of normal distribution using the Shapiro-Wilk test for normality. The mean and standard deviation were calculated. Means were compared between groups using 2-tailed Student’s t-tests (α=0.05).

The mean and standard deviation were calculated. Means were compared between groups using 2-tailed Student’s t-tests (α=0.05). The results are shown in Table 1.Table 1: Results of the tensile testing of the cortical bone analog material.Modulus of Elasticity (GPa)Fiber Volume Fraction0.0 0.18 0.30 0.39Average3.091 5.954 7.890 10.607Standard Deviation 0.271 0.410 0.854 1.33195% Confidence Interval 2.839 - 3.342 5.574 - 6.333 7.100 – 8.681 9.377 – 11.83All of the samples with glass fiber had a significantly higher modulus of elasticitycompared to the samples made without any fiber (p<0.001). The samples with a fiber volume fraction of 0.30 had a modulus of elasticity that was significantly higher than the samples with a fiber volume fraction of 0.18 (p<0.001). The samples with a fiber volume fraction of 0.39 differed statistically from the samples with a fiber volume fraction of 0.30 (p<0.001).

The samples with a fiber volume fraction of 0.39 differed statistically from the samples with a fiber volume fraction of 0.30 (p<0.001). The high viscosity in the samples with 0.39 fiber volume caused some difficulty when mixing, degassing and filling the mold which may have led to some the higher standard deviation. In general, the measured tensile moduli fell within the expected range of cadaveric properties.2.2 Cancellous Bone Analog DevelopmentCancellous bone is commonly found in the epiphysis of long bones. The structure ofcancellous bone is very similar to that of open celled foam, and unlike the cortical bone analog, there is widespread documentation in the literature regarding material formulations for cancellous bone analog materials. Typically, rigid polyurethane foam materials are used as a model for cancellous bone. ASTM F1839 provides the specifications for preparing polyurethane samples for use in testing orthopedic devices.

Typically, rigid polyurethane foam materials are used as a model for cancellous bone. ASTM F1839 provides the specifications for preparing polyurethane samples for use in testing orthopedic devices. This material is easily mixed and poured/injected into molds, cures quickly, and exhibits material properties with much less variability compared to cancellous bone specimens obtained from cadavers22. By varying the ratio of isocyanate to resin in the mixture, Szivek showed that the mechanical properties of polyurethane foams could be adjusted to match the mechanical properties of a wide range of cancellous bone materials.The tensile elastic modulus of cancellous bone varies depending upon the location in thebody, the presence of disease, age, etc. In a book reviewing of multiple studies on the mechanical properties of bone, Yuehuei and Draughn found that the tensile modulus of elasticity of cancellous bone ranged from 69 MPa to 349 MPa (10 ksi to 50.6 ksi).

Tensile testing on a polyurethane cancellous bone analog material was carried out in accordance with ASTM D638- 08. Polyurethane foam (16lb/ft3, U.S. Composites, West Palm Beach, FL) was used to prepare the cancellous tensile specimens, it was mixed in a ratio of 1:1 (by volume) and poured into the same silicone rubber mold that was used to fabricate the cortical bone tensile specimens The polyurethane cured for 1 hour at 21ºC (70ºF). Moisture in the air may affect the properties of the foam during curing, therefore the relative humidity was recorded at 42%. Tensile testing was carried out at a loading rate of 1.27 mm/min (0.05 in/min). The average tensile modulus of elasticity was 305MPa (44.3 ksi) with standard deviation of 30.06 Mpa (4.36 ksi).Figure 3: Images showing the microstructure of the polyurethane cancellous bone analog (left) and cancellous bone sectioned from the distal epiphysis of a canine radius (right).

Scale bar indicates 200µm.3.0 Design and Manufacture of a Patient-Specific Bone Analog ModelIn the previous sections the materials have been developed that mimic some of theintensive mechanical properties of cortical and cancellous bone. This section describes rapid prototyping, and rapid tooling methods used to combine these materials into a patient specific geometry that also mimics the extensive mechanical properties of the whole bone of a given patient. The mechanical properties of the whole analog models of canine radii were compared to a group of cadaver bone in a separate experiment14.3.1 Design and Fabrication of Rapid Tooling for Bone Analog ModelsNine adult dogs weighing more than 30 kg (66.14 lbs) were collected after euthanasia. The dogs were euthanatized at a local animal shelter for reasons unrelated to this study. Age, breed, gender, and method of euthanasia were recorded.

The dogs were euthanatized at a local animal shelter for reasons unrelated to this study. Age, breed, gender, and method of euthanasia were recorded. The thoracic limbs were disarticulated at the glenohumeral joint, wrapped in saline solution (0.9% NaCl) moistened gauze, sealed in a plastic bag, and frozen at -20°C (-4°F). The limbs were later thawed at room temperature and computed tomography imaging was performed using a helical CT (Siemens SOMATOM Sensation 16-slice configuration, Siemens Medical Solutions, Malvern, Penn) with 512 x 512 resolution and 0° gantry tilt. The CT images were retro-reconstructed into 1 mm (.039 inch) slices with a pixel size of 0.727 mm (0.029 inch). Materialise Mimics software (Mimics version 12.1, Materialise, Belgium) was used to isolate and reconstruct the three-dimensional geometry each specimen.

Materialise Mimics software (Mimics version 12.1, Materialise, Belgium) was used to isolate and reconstruct the three-dimensional geometry each specimen. Built in NURBS algorithms were used to generate a closed surface model for the cancellous and cortical regions of the bone (the cancellous portion has been extended to include the medullary canal for ease of manufacture).These models were then imported into Solidworks 2009 3-D CAD modeling software in order to design the mold and mold parting lines, add geometric features for mold filling, venting, as well as features that facilitate the three dimensional alignment of the cancellous bone analog insert within the cortical bone mold. The models were then converted into stereolithography (.stl) format and fabricated in plastic using rapid prototyping technology (Objet Alaris 30). From the physical models, polyurethane rubber molds were fabricated.

From the physical models, polyurethane rubber molds were fabricated. For each mold, the plastic model was placed into an adjustable acrylic plastic mold box, the parting line was marked on the models. The models were then coated with mold release agent (Ease-release 200, Smooth-on Inc.). Room temperature vulcanizing rubber (Reoflex 60, Smooth-on Inc.) was mixed in a 1:1 ratio (by weight) and degassed in a vacuum chamber, and then poured into each half of the mold and allowed to cure overnight. Figure 4 shows several images from the process. The rubber used for the molds was a relatively high durometer (60 Shore A) in order to minimize mold deformation during injection molding, yet still provide some flexibility to facilitate the demolding of parts with slight undercuts. Molds were post cured for 4 hours at 75ºC (167ºF).

Molds were post cured for 4 hours at 75ºC (167ºF). A clamp was fabricated using stainless steel strapping in order to hold the molds halves closed during injection/filling.Figure 4: Images showing: a screenshot of CT data (a); images of the cortical and cancellous bone 3-D models with the filling sprue (1), air vents (2) and 3-D geometric locating features (3,4) (b); finished plastic models of the cortical and cancellous models (c); and polyurethane rubber mold (d).3.2 Fabrication of Bone Analog ModelsTo fabricate the cancellous bone inserts, the mold halves were coated with mold release agent and then closed together using the custom clamps. Ten milliliters (0.610 in3) of the 16 lb/in3 density polyurethane were combined and mixed for 30 seconds, and then poured into the cancellous insert mold. Initial testing indicated that the polyurethane material would deform and break when subjected to the forces from injection molding the fiberglass cortical bone analog.

Initial testing indicated that the polyurethane material would deform and break when subjected to the forces from injection molding the fiberglass cortical bone analog. For this reason, a 3.175 mm (0.125 inch) brass rod was placed in the cancellous bone mold, oriented axially. This rod was easily removed from the final bone analog, and leaving the medullary canal hollow. The mold was capped, and the polyurethane mixture was allowed to cure for 25 minutes at 21ºC (70ºF) and 42% relative humidity. After curing, the models were removed from the mold and rinsed with acetone solution to remove the mold release agent. To prevent the open pockets on the surface of the foam models from creating bubbles in the final product, the surface was coated with a thin layer of epoxy resin (Loctite 5 min epoxy, Henkel, Düsseldorf, Germany).

In preliminary tests it was noted that the exothermic curing process of the epoxy/fiberglass generated enough heat to break down the polyurethane material, creating large bubbles. Therefore, the surface of the cancellous bone insert was coated with a layer of high temperature ceramic coating (VHT Ceramic Coating, Sherwin-Williams) that eliminated the formation of bubbles. The coating was allowed to dry for one hour.Mold release agent was then used to coat the surfaces of the cortical bone mold. Usingthe built-in geometric locating features, the cancellous insert was precisely located within the cortical bone mold, the mold was closed using the aforementioned clamps. 200 g (0.441 lbs ) of epoxy resin and short fiber glass in the ratios described in section 2 were mixed under vacuum using a custom-built mixer within a vacuum chamber for 10 minutes. While still under vacuum, the epoxy-fiberglass mixture was drawn into a 60 cc (3.66 in3) syringe.

While still under vacuum, the epoxy-fiberglass mixture was drawn into a 60 cc (3.66 in3) syringe. The mold and syringe were then loaded into a custom-built injection molding machine that introduced the resin into the mold with a pressure of 68.94 kPa (10 psi). Figure 5 illustrates this process.Figure 5: Photographs showing several steps of analog bone fabrication process; molding of the polyurethane cancellous bone analog insert (a), placement of the ceramic coated cancellous bone analog insert within the cortical bone analog mold (b) mixing of the epoxy-fiberglass mixture under vacuum(c), and injection of epoxy-fiberglass mixture into the cortical bone analog mold (d).Immediately after mold filling had occurred (verified visually by the presence of theepoxy-fiberglass mixture in the mold air vents), the mold was placed inside of a custom-built pressure vessel at 482.6 kPa (70 psi) for one hour until the epoxy cured.

This reduced the volume of any bubbles that formed either as a result of the epoxy curing process, or from the interaction between the epoxy and the polyurethane cancellous bone analog insert. The samples were then post cured in a convection oven at 75ºC (167ºF) for one hour before being removed from the mold.One of the key parameters for the prediction of the modulus of elasticity of the short fiber composite bone analog is the orientation of the fibers within the composite. To evaluate the fiber orientation within the composite bone analogs, special bone samples were fabricated in which the epoxy matrix was dyed dark blue (Smooth-on Inc.). Cross sections were cut from the bone analog models which were progressively ground and then polished (120-2000 grit paper) using standard methods (ASTM E2015-04). The microstructure of the composite was then digitally photographed at a 350x magnification. LabviewTM software was used to threshold and filter noisy data points.

The microstructure of the composite was then digitally photographed at a 350x magnification. LabviewTM software was used to threshold and filter noisy data points. Labview’s built-in ellipse fitting algorithm was used to maculate the orientation, the center, the major axis (a) and the minor axis (b) for each fiber cross-section. From this a two-parameter fiber orientation distribution ψ= [Ф, θ] was measured; where θ is the in-plane fiber orientation and, Ф is the out-of-plane fiber orientation (cos-1 b/a).Despite the ambiguity of the in-plane orientation (i.e θ or θ+π) a close estimate of fiberorientation can be obtained26. The analysis indicates that the fibers are randomly oriented in plane, but are relatively aligned with the longitudinal axis of the bone (out-of-plane). This is a relatively common effect seen in short-fiber reinforced injection moldings.

This is a relatively common effect seen in short-fiber reinforced injection moldings. Figure 6 shows a photograph of several completed bone analog models, the microstructure of the cortical bone analog, and histograms showing the in-plane and out-of-plane fiber orientations.Figure 6: Photograph of several custom analog radii prepared for mechanical testing (a), the microstructure of the analog bone model (b) and histograms showing the in-plane and out-of-plane fiber orientations (c).4.0 Preliminary Evaluation of the Potential for Direct Fabrication of Patient-Specific Analog Bone Models Using Additive Manufacturing MethodsIn the previous sections, computed tomography, 3-D modeling, additive manufacturingand rapid tooling were utilized to quickly manufacture replicas of a patient-specific bone analog model with intensive mechanical properties similar to that of the original bone. The mechanical properties of these models have less variability than cadaveric specimens.

The mechanical properties of these models have less variability than cadaveric specimens. However, the process requires a significant investment in time and effort (which translates into cost). This includes the scanning of the patient, design of the mold in a CAD system, fabrication of the cortical and cancellous mold plugs, fabrication of the mold, and then finally the fabrication of the analog bone models. One potential solution would be to fabricate the analog bone models directly from CT data using a new form of additive manufacturing that incorporates a fiber reinforced material similar to the epoxy resin used in this study. Several studies have examined the addition of fiber reinforcements to traditional additive manufacturing materials to improve the mechanical properties26.

Several studies have examined the addition of fiber reinforcements to traditional additive manufacturing materials to improve the mechanical properties26. As an initial examination of this possibility, seven tensile specimens of each fiber fraction used in section 2 of this study; 0%, 33% , 50% , and 60% fraction by weight (0%, 18%, 30%, and 39% by volume) were prepared. The matrix material used was DSM Somos Watershed 11120 ultraviolet curing resin (DSM Somos Elgin, Il). This material has a tensile modulus of elasticity 2.59 Gpa (376.2 ksi). The resin and fiber were mixed under vacuum for 10 minutes and then poured into an open silicone rubber mold with the same geometry as the tensile specimens in section 2 (conforming to ASTM D638-08). The mold was placed into an ultraviolet chamber (PCA250 3D Systems) and cured for 12 hours, after which the samples were flipped and then cured for another 12 hours.

The mold was placed into an ultraviolet chamber (PCA250 3D Systems) and cured for 12 hours, after which the samples were flipped and then cured for another 12 hours. In addition samples of the same geometry were fabricated with the same resin (0% fiber) but using the stereolithography process (SLA 250, 3D Systems Rock Hill, SC). The samples with 39% fiber volume fraction did not cure properly, and were therefore excluded from the test.Tensile testing was carried out with the same methodology and equipment as previously described. The loading rate was 1.27 mm/min (0.05 in/min). The modulus of elasticity of each sample was calculated from the linear portion of the stress-strain plot. After confirmation of normal distribution using the Shapiro-Wilk test for normality, the mean and standard deviation were calculated for each test and reported. Means were then compared between groups using 2- tailed Student’s t-tests.

Means were then compared between groups using 2- tailed Student’s t-tests. The results are shown in Table 2.Table 2: Results of the tensile testing of the cortical bone analog material.Modulus of Elasticity (GPa)Fiber Volume Fraction0.0 0.0 (SLA fabricated) 0.18 0.30 0.39Average 2.594 2.752 3.845 6.116 NAStandard Deviation 0.089 0.054 0.270 0.128 NA95% Confidence Interval 2.372 – 2.816 2.617 – 2.886 3.174 – 4.516 5.797 – 6.435 NAThere was no significant difference between the 0% fiber tensile specimens madeusing the SLA additive manufacturing process and the 0% fiber tensile samples cured in the mold (p=0.059). The samples with 30% fiber (by volume) had a higher modulus than any of the other groups (p<0.001), the 18% fiber samples (by volume) had a modulus of elasticity that was greater than the 0% fiber samples (p<0.001) and less than the 30% fiber samples (p<0.001).

The modulus of elasticity of the cortical bone analog samples made with UV curing resin and fiberglass were compared to that of the cortical bone analog made from epoxy resin and fiberglass. With the exception of the 39% (by weight) fiber, which was excluded from testing, the samples made with the epoxy and fiberglass had a higher modulus of elasticity compared to the samples made using UV resin and fiberglass (p<0.001).5.0 Discussion and ConclusionsThis paper has described a new platform upon which the mechanical properties of custombiomedical implants can be evaluated. The cortical bone and cancellous bone analog materials described herein have been shown to exhibit a tensile modulus of elasticity, a key intensive material property, within the expected range of cadaveric specimens and with a significant reduction in variability.

This is advantageous on several levels, particularly because it allows the reduction in the required sample size while also providing the ability to evaluate multiple implant parameters simultaneously. Computed tomography combined with traditional, plastic-based, additive manufacturing facilitates the fabrication of rapid tooling for patient-specific bone geometries which will support the mechanical testing and evaluation of custom implants and devices.The development of a short-fiber reinforced composite cortical bone analog material hasalso been translated to matrix materials typically used in rapid prototyping processes with promising results. This indicated that with the appropriate equipment modifications, the direct fabrication of custom, patient-specific bone analog models is possible.6.0 References1. Cansizoglu O., Harrysson O. L. A., Cormier D. R., West II H.

sign and Optimization of a High Temperature Microheater for Inkjet DepositionAustin VanHorn, Wenchao Zhou The AM3 Lab, Department of Mechanical Engineering, University of Arkansas at Fayetteville Fayetteville, AR, United States of America; email: zhouw@uark.eduABSTRACTInkjet deposition has become a promising additive manufacturing technique due to its fast printing speed, scalability, wide choice of materials, and compatibility for multi-material printing. Among many different inkjet techniques, thermal inkjet, led by Hewlett-Packard and Canon, is the most successful inkjet technique that uses a microheater to produce a pressure pulse for ejecting droplets by vaporizing the ink materials in a timespan of microseconds. Thermal inkjet has been widely adopted in many commercial 3D inkjet printers (e.g., 3D Systems ProJet X60 series) due to its low cost, high resolution, and easy operation.

Thermal inkjet has been widely adopted in many commercial 3D inkjet printers (e.g., 3D Systems ProJet X60 series) due to its low cost, high resolution, and easy operation. However, the viscosity of the printable materials has been limited to less than 40cP due to insufficient energy provided inside the nozzle to overcome the viscous dissipation of energy. This paper presents a study on the design and optimization of a high temperature microheater with a target heating temperature of more than 600˚C (compared to ~300 ˚C for current printhead) to increase the energy supply to the nozzle.

The benefits are fourfold: 1) higher temperature will lead to faster vaporization of ink and thus higher jetting frequency and print speed; 2) higher temperature will make it possible for jetting materials with higher boiling points; 3) higher temperature will reduce the viscosity of the ink and thus the viscous dissipation of energy; 4) higher energy supply will increase the magnitude of the pressure pulse for printing more viscous materials. In this paper, a high temperature microheater was designed with the following objectives: to reduce thermal stress in heaters, and to minimize uneven heat distribution. A literature survey was first conducted on design, fabrication, and operation of thin-film resistive microheaters. A multiphysics numerical model was then developed to simulate electrical, thermal, and mechanical responses of the microheater. The model was validated by comparison to experimental data and existing models obtained from literature.

The model was validated by comparison to experimental data and existing models obtained from literature. With proper parameterization of the design geometry, the geometry of the microheater is optimized using a particle swarm optimization method. Results show the optimized high temperature microheater successfully operates at temperatures in excess of 600˚C. The design optimization enabled better characteristics for even heat distribution and minimizing stress. The design approach can serve as a fundamental means of design optimization for microheaters.I. INTRODUCTION Thermal inkjet has become a promising additive manufacturing technology due to its printing speed, low cost to fabricate, and fine printing resolution. It uses a microheater to produce a pressure pulse for ejecting droplets by vaporizing the ink materials in a timespan of microseconds.

It uses a microheater to produce a pressure pulse for ejecting droplets by vaporizing the ink materials in a timespan of microseconds. The microheater in the current thermal inkjet printhead typically operates at ~300°C [1] and the viscosity of printable materials is limited up to ~40cP [2]. Increasing the operating temperature of the microheater may offer multiple benefits, including higher jetting frequency and wider choice of materials with high viscosity and boiling points. Extensive research has been conducted on microheater development, such as fabrication, material properties, temperature limits, failure modes, and sustainability [3-11]. Progress has also been made on development of high temperature microheaters that can reach above 600 ˚C. Simulation has been used to predict the behaviors and optimization has been performed through iterativesimulations and experiments [11, 12].

Simulation has been used to predict the behaviors and optimization has been performed through iterativesimulations and experiments [11, 12]. However, little research has been reported on systematic approach for optimization of microheater geometry. In this paper, a constrained particle swarm optimization method, in combination with finite element simulations, is developed for optimization of a high temperature microheater that can operate at 600 ˚C, with design objectives to minimize uneven temperature distribution and thermal stress under certain design constraints. This study first presents a literature review on the current status of microheater design. It then proceeds into the development of a numerical model for simulating microheater operation. The numerical model is validated with experimental data from literature [9, 10].

It then proceeds into the development of a numerical model for simulating microheater operation. The numerical model is validated with experimental data from literature [9, 10]. With proper parameterization of the microheater geometric design, the validated numerical model is used to evaluate the designs with respect to the design objectives and a particle swarm optimization method is developed to find the optimal microheater geometry in a constrained design space.II. LITERATURE REVIEWMany research efforts were reported on thin-film resistive heaters (microheaters). This section is to provide a review of the state of the art of microheater technology and identify the research gap.A. History and MotivationJames Prescott Joule initiated the first research conducted in resistive heating in 1841 which was later followed Hienrich Lenz in 1842.

History and MotivationJames Prescott Joule initiated the first research conducted in resistive heating in 1841 which was later followed Hienrich Lenz in 1842. This method of heating did not get accepted by the Royal Society in London until 1849 when Michael Farraday sponsored the findings, which soon after led to James Joule’s acceptance in the Royal Society [13]. Since then resistive heaters have found their way into many designs. One such application of resistive heating is thermal inkjet deposition. HP first introduced thermal inkjet technology in 1984 with their dot-matrix printer which miniaturized inkjet technology with higher quality, lower power, and quieter operation [14]. Current microheaters in thermal inkjet printers typically operate at ~300°C [1], which limits the ability to print materials with higher viscosity and higher boiling points.

Current microheaters in thermal inkjet printers typically operate at ~300°C [1], which limits the ability to print materials with higher viscosity and higher boiling points. Higher temperatures will increase the pressure pulse from thermal inkjet printers and create quicker vaporization of materials therefore higher frequency for droplet generation.B. Microheater DesignA microheater is a sophisticated MEMS device and its operation involves multiple physics, (electrical, mechanical, and thermal) as well as materials and geometric designs. Classical design of a microheater consist of four elements [15]: 1.) conductive pad, 2.) resistive heating element, 3.) substrate, 4.) oxidation passivation layers. A typical configuration is shown in Figure 1.Figure 1-Schematic of the Four Basic Elements of a MicroheaterThe actual configuration of microheaters however, is ultimately based on the parameters and objectives trying to be reached with a particular design.

One of the parameters establishedfor our design is to reach temperatures of 600°C. Much research has been done on microheaters capable of reaching temperatures of 600°C, which involves materials, fabrication, modeling, and design. 1) MaterialsMany materials have been researched in literature for the four components used in a typical microheater configuration. The substrate provides the foundation for the microheater. Typical materials used are glass [10], silicon [8], ceramics [9], and stainless steel [16]. Silicon and ceramics typically can handle higher ranges of temperatures than steel and glass [16]. Ceramics like alumina typically have a high thermal conductivity creating high power consumption [9, 17]. However, Low Temperature Cofired Ceramics (LTCC) are a composite ceramic/glass material typically offering much lower thermal conductivity enabling power savings [9].

However, Low Temperature Cofired Ceramics (LTCC) are a composite ceramic/glass material typically offering much lower thermal conductivity enabling power savings [9]. Stainless steel has higher thermal conductivity, but is often chosen because it is inert, robust, and has many well defined microfabrication techniques [16]. Silicon is a frequently chosen material used for substrates due to its ability to operate at 1000ºC temperatures for prolonged periods of time and high thermal conductivity enabling quick thermal response times [8].The resistive layer is where the main source of resistive heating occurs. Resistive materials that can operate at temperatures above 600°C are needed for high temperature microheaters. Some of the studied materials include Platinum (Pt) [9, 18, 19], Titanium (Ti) [10, 15], SnO2:Sb (Sb doped in SnO2) [8], Si:B [16], Si:P [16], Tungsten [3], and Molybdenum [5, 20]. The maximum operating temperatures of Pt and Ti are 600°C [16] and 700°C [15].

The maximum operating temperatures of Pt and Ti are 600°C [16] and 700°C [15]. SnO2:Sb and poly-silicon materials have temperature limits up to 1000°C, but require high voltage input due to their high resistivity [8]. Tungsten also has a high temperature operation range of 1200°C, however tungsten is only stable in an oxygen-enriched environment up to 400°C [16]. Molybdenum like Tungsten suffers significant oxidation at temperatures of 300°C, which becomes volatile at 700°C, but has capabilities of reaching temperatures of 850°C in a shielded environment [5]. Some material properties in the resistive layer also change with geometry, but the impacts of geometry on resistivity have been shown to be minimal within thin films [4, 11]. Sondheimer theory states that the change in the thickness of the material results in minimal difference in material properties unless the material thickness is less than 8% of its mean-free- path for thin-film metals.

This has also been demonstrated experimentally in [21]. Common materials used for oxidation layers are , TiO2, and Al2O3/Pt/Ta [16]. These ceramic, silicon, and metal oxide materials have been known to cause high stresses in microheaters [22]. These layers provide many uses in microheaters. They are typically used as insulators to create power efficient microheaters [23]. They’re also used for electrical insulation and as a protecting layer [24].Another material, not shown in the typical configuration, is adhesive materials. Use of adhesives to bond resistors to substrates or oxidative layers is a common practice. Typical materials include Ti, Ta, and Zr [18]. However, at temperatures over 700°C Ti normally will suffer from adhesion issues failing due to the high shear stresses causing the materials to peel off from the resistor or the substrate [25].

However, at temperatures over 700°C Ti normally will suffer from adhesion issues failing due to the high shear stresses causing the materials to peel off from the resistor or the substrate [25]. The effects of stress have been minimized by using intermediate layers such as Silicon Nitride so that the thermal stress between layers isn’t as substantial [26].2) ModelingModeling of microheaters has been used extensively for microheater design. Modeling includes the electrical, thermal, and mechanical characterization of microheaters. An important aspect to the electrical and thermal characterization is the change in materials properties as temperatures are varied. Experiments have been performed to determine the values of materials properties at different temperatures [10]. Resistivity is a material property of particular importance due to its critical role in Joule heating.

Resistivity is a material property of particular importance due to its critical role in Joule heating. Materials have a tendency to increase resistivity as temperature increases [27].Prior research has been performed by experimentation evaluating the effects of heat transfer on microheaters. Microheaters are affected by three types of heat transfer which are conduction, convection, and radiation [6]. It was shown that the effects of conduction and convection can have a significant impact on the heat transfer while radiation is negligible for Ti and Pt materials at temperatures of less than ~700°C [6, 10, 19].Modeling has also been an essential part of design for microheaters. Modeling has been used to select materials by evaluating the maximum temperature and power savings of microheaters with different insulating layers [9]. In two other studies, modeling was used to study the effects of microheater geometry to improve the evenness of heat distribution.

In two other studies, modeling was used to study the effects of microheater geometry to improve the evenness of heat distribution. Unique geometries were assessed in [11]. The second study used a spiral pattern, and the effects of the geometric parameters were investigated, such as the filament width and spacing between the filament [12].3) DesignMany noteworthy design aspects, such as material properties [10], electro-thermal response [9], heat distribution [11], and geometry [12] were found from literature in previous research. K.L. Zhang et al. evaluates the material properties of a Ti/Au microheater. In this study, the design impacts of the materials thermal conductivity, resistivity, and emissivity all have significant impacts on the design. With increase in temperature, the emissivity and resistivity of materials increases while the thermal conductivity decreases [10].

With increase in temperature, the emissivity and resistivity of materials increases while the thermal conductivity decreases [10]. This study also made a comparison of Ti/Au resistors to Pt resistors showing the impacts of thermal expansion and high thermal stress was the typical mode of failure in Ti microheater design. Even heat distribution has been shown to provide power savings by being more efficient and also creates better responses times in microheaters [11]. Thermal response time reached 2ms for high temperature designs of 600°C for Tungsten microheaters with low power consumption of 12mW [3]. A time response of 1ms was achieved in a Pt/Ti heater reaching a temperature of 400°C using only 9mW of power [28].C. Particle Swarm OptimizationParticle swarm was initially a method of predicting social behavior in animals [29].

Particle Swarm OptimizationParticle swarm was initially a method of predicting social behavior in animals [29]. This method has been adopted for an extensive amount of applications [30], such as array failure correction for antennas, predictive and tuning controls, design optimization, etc. Particle swarm optimization is an advantageous method. It enables optimization of parameters within a discontinuous problem and with much less computational cost than other optimization method. This is enabled by the method of stochastically searching for the optimal solution [31]. The method evaluates the fitness of a current set of solutions and identifies new design parameters based on a fitness evaluation for minimizing an objective function. Each solution generated moves the parameters closer to the optimal solution at which the numerical problem reachesconvergence [32].

Each solution generated moves the parameters closer to the optimal solution at which the numerical problem reachesconvergence [32]. One recent publication utilized particle swarm optimization for design of microheaters to create power savings in microheaters [33]. This application uses basic equations of heat transfer to find the optimal design parameters. The parameters used were the oxide layer thickness, temperature, and active area. Active area is the area of the heat source. No attention was given to optimize for stress nor was the actual geometry of the microheater changed. They also did not consider any constraints, or utilize a finite element model.III. MODELING SETUP AND VALIDATIONModeling of a microheater involves multiple coupling physics. The physical phenomenon experienced in a microheater is that of Joule heating. Joule heating consists of electric currents and heat transfer.

The physical phenomenon experienced in a microheater is that of Joule heating. Joule heating consists of electric currents and heat transfer. In this simulation the structural mechanics involved were also considered which were coupled with heat transfer by thermal expansion.A. Numerical ModelingIn this section, a model for microheater operation is developed. This section walks through the governing equations used to predict the electrical, thermal, and mechanical response of a microheater.1) Current conservationJoule heating is created by running current through a resistive material.

The constitutiverelationship for the resistive material is described by Ohm’s Law:Equation 1J is the current density, E is the electric field strength, and σ is the electrical conductivity:[ ( )]Equation 2 where represents the electrical resistivity, α is the temperature coefficient of resistivity with the assumption that the resistivity changes linearly with temperature, and ( ) is the change in temperature of the resistive material. Based on charge conservation and equation of continuity, the electric potential throughout the microheater can be described by:( ) Equation 3 where is the layer thickness, V is the electrical potential applied, and denotes the gradient operator in the tangential direction of the electric field. Resistive heating is induced from the resistivity in the circuit opposing electron flow. This opposition creates heat that can be shown by:| | Equation 4 where Q is the heat being generated.

Resistive heating is induced from the resistivity in the circuit opposing electron flow. This opposition creates heat that can be shown by:| | Equation 4 where Q is the heat being generated. The equations are coupled together based on the governing equations dependency on other physical phenomenon.2) Heat TransferHeat transfer follows the law of thermodynamics. In microheaters, heat produced by Joule heating is primarily transferred away by conduction and convection for temperature below700 ˚C. Since only these forms of heat transfer are present, the first law of thermodynamics ultimately reduces to the governing equation for heat transfer:() ( )Equation 5 The variables quantifying this equation are the density , specific heat , thermal conductivity , convection coefficient h, change in temperature , and heat source Q.

represents the change of temperature with respect to time and is the gradient operator.3) Structural MechanicsStress is induced in microheaters as a result of thermal expansion when materialsexperience a rise in temperature. The governing equation for the structural mechanics is:( )Equation 6where s is the stress tensor, is material displacement, and Fv is the volume force vector.Computational cost of modeling microheaters can be extensive due to the large ratio of the thickness of the resistive material layer (typically ~200 nm) to that of the substrate (~ 25 µm), which could cause significant meshing issues. One way to simplify is by using the shell theory. Structural mechanics shell theory is different from three dimensional structural mechanics in that the models are formulated by using the Lagrange continuum mechanics with mixed interpolation of tensor components. This reduces computations to consider only plane stress since other stresses will have minimal impact.

This reduces computations to consider only plane stress since other stresses will have minimal impact. The mechanical stress is formulated as:( )( )&Equation 7Equation 8 where σ is the in plane stress, ( ) is the volume moment vector, z is the vertical direction, and d is the overall thickness of the shell. The stresses induced in the design are all resultants of thermal expansion which is shown from the thermal strain as:( ) & Equation 9( )Equation 10 where corresponds to the strain subject to the bending stress in the shell from thermal expansion, is the strain in the entire domain from thermal expansion, and is the thermal coefficient of expansion. The thermal expansion can be coupled to Equation 6 and Equation 7 using the stress-strain relationship of materials.B. Validation of Numerical ModelThe numerical model was implemented using COMSOL 5.1.

Validation of Numerical ModelThe numerical model was implemented using COMSOL 5.1. A multiphysics model was developed for a Pt miroheater on an alumina composite substrate by coupling three physicsinvolved (i.e., electrical, mechanical, and thermal). This study evaluated the temperature change in the microheater and substrate.

This study evaluated the temperature change in the microheater and substrate. To form a solution the numerical model was evaluated transiently.Figure 2-Heater geometry and Boundary ConditionsTable 1-Material Properties of for the 1st Numerical Model Validation SummaryPtAl2O3Specific Heat (J/kg*K)133800Thermal conductivity (W/m*K)71.625Coefficient of Thermal Expansion (1/K)Density (kg/m^3)8.80E-06 214506.50E-06 3780Temperature Coefficient of Resistance (1/K)3.93E-03Resistivity (Ω*m)8.33E-07Young's Modulus (Pa)1.68E+114.00E+11Poisson's RatioFunction0.38 resistor0.22 substrateThickness (µm)0.61000Figure 2 and Table 1 show the initial setup of the numerical model, where the shell boundaries of the Pt resistor were centrally located on the substrate. The open boundary condition was placed on the boundary where the domain would be continuous.

The open boundary condition was placed on the boundary where the domain would be continuous. This assumes | so the temperatures of the un-modeled part have continuous temperature change at the boundary. The convective heat loss coefficient applied to all other sides was 5W/m*K, which has been commonly used in literature [9, 10]. Table 1 shows the material properties.The model was validated by comparing the simulation results with the experimental data obtained from literature [9], which are shown in Figure 3. The change of temperature at the center of the microheater after 200 ms with respect to input voltage was compared.

The change of temperature at the center of the microheater after 200 ms with respect to input voltage was compared. As we can see, the results have an excellent agreement between our numerical simulations and the experimental data from literature.Figure 3-Change of Microheater Vemperature with Voltage Input of Numerical Simulation Compared with Experimental DataA second validation was also performed on the transient response of the microheater over time because of the importance of transient behavior of a microheater. Simulations of a Ti/Au resistor on a 7740-Pyrex glass substrate were performed to compare with experimental results and other modeling results from literature [10].To simplify this model the conductor layer composed of Au/Ti was modeled as two resistors in parallel. This was assumed due to the parallel orientation in which the materials were stacked and current was supplied through the conductive layer. The convection coefficient used in this model was 100W/m*K [10].

A 3V potential difference was placed on the resistor, and then the temperature at the hottest point over the entire microheater configuration was extracted at each time step. The transient temperature change over time was compared to previous experimental and modeling results shown in Figure 5 [10]. From comparison, we can see a good agreement. The model being validated agrees more closely with the experimental data as it approaches a steady state than the results from the previous model [10].Figure 5-Comparison of Numerical Model to Previous Research for 2nd ValidationIV. OPTIMIZATIONIn this section, a particle swarm optimization method was developed to optimize the microheater design in order to minimize its thermal stress and even heat distribution during operation. An initial design was first established and design parameters were identified.

An initial design was first established and design parameters were identified. The validated numerical model was then employed to evaluate different designs with respect to the optimization objective (i.e., minimal stress, even heat distribution). A particle swarm optimization method was then developed to search for the optimal design in the design space.A. Initial Design and Design VariablesBased on our literature review, an initial design of the microheater was chosen as follows. An Ultra-thin Glass (type AF 32*eco from Schott Inc.) was used as the substrate to support the resistive filament and conductive pads. This material was chosen due to its low thermal making it power efficient [10]. It also has a transition temperature of 717ºC making it capable of reaching the desired operating temperature range before reaching the transition temperature. Titanium was chosen as the resistive material because of its capability of reaching temperatures in excess of 600°C.

Titanium was chosen as the resistive material because of its capability of reaching temperatures in excess of 600°C. Its failure from thermal stress is often the limiting factor, but not until temperatures are in excess of 700°C. Gold was chosen due to its high electrical conductivity. Gold also has low coefficient of thermal expansion, as compared to other popular conductors like silver, which induces less stress in the design.

It was demonstrated to be capable of creating a minimal temperature deviation across the design space. Figure 9 shows the initial design. The outer radius determines the size of the microheater and was chosen to be 50um for this study. The initial values of the rest of the design parameters are listed in Table 4.Figure 6-Design Parameters Describing Topology of the ResistorTable 4-Initial Values of the Design ParametersParameter Value Unit Um Fil_wid Um Outer_s Um Vert_cut Um Horz_cut Um Inner_s Nm Fil_thick12.5 27.5 25 12.5 10 200Description Filament Width Outer Spacing Vertical Cut Horizontal Cut Inner Spacing Filament ThicknessB. Particle Swarm OptimizationWith the six design parameters chosen, we needed to determine the optimal combination of the design parameters in the six-dimensional design space. Because it was impossible to test every design in the design space, an efficient method was needed to search for the optimal design.

Because it was impossible to test every design in the design space, an efficient method was needed to search for the optimal design. Particle swarm optimization has established itself as a popular method of design optimization. This method stochastically searches for an optimal solution within a specified design space by changing the design parameters. We developed a unique method for constraining particle swarm algorithms as well as using complex models within the algorithms. This is a new concept to microheater design. In this section, we will use constrained particle swarm optimization to search for optimal topology solutions using the variables initially established in the geometry.Particle swarm optimization randomly searches for solutions to problems within a design space. In this algorithm, each particle is a design with a set of design parameters (i.e., the six design parameters in Table 4.

In this algorithm, each particle is a design with a set of design parameters (i.e., the six design parameters in Table 4. The lower bound and upper bound of the design parameters determine the boundaries of the design space. Each particle has its own unique position and velocity in the design space. The velocity is determined by the particle’s personal best and the global best of all particles and used to updates the particle’s current position based on:( ) ( ) ( ) ( ) Equation 11Where c1=c=2 and are arbitrary constants, pbest is the personal best of the particle, rand is a random number from [0,1], gbest is the overall best is the algorithm, present is the present solution, and vi is the current velocity.The validated finite element model was used to evaluate the fitness of each particle with respect to the design objectives. The algorithm randomly generates initial particles and moves the particles in the design space based on their velocities.

The algorithm randomly generates initial particles and moves the particles in the design space based on their velocities. This stochastic search enables convergence of optimization algorithms for complex problems such as a microheater design. Certain designs are undesirable for the microheater, which were accounted for by placing constraints on the particles. One constraint was that we want the total resistance of the microheater to be above 100 Ω so that the heat concentrates on the microheater instead of on the external circuit. Another constraint was used to ensure geometric continuity in the design.

Results and DiscussionBoundaries of each design parameter are defined in Table 5. The first objective function was to minimize the maximum stress over the entire microheater. Only thermal stresses in the substrate are accounted for in this design. Stress is sometimes induced in the devices from fabrication methods and exist in other parts of the microheater (i.e. filament, adhesive interface), but are not taken into account here. The second objective function was to create even heat distribution by minimizing the standard deviation of temperature within 121 evenly distributed sample points over the entire microheater. The amount of sample points used to evaluate the model was determined to be sufficient in this study. Even heat distribution was accomplished within the sample space by optimizing variables input to the optimization algorithm. For the simulations, power was supplied at 1W for 1 ms. The simulation creates a temperature of ~600ºC within the design for both objectives.

For the simulations, power was supplied at 1W for 1 ms. The simulation creates a temperature of ~600ºC within the design for both objectives. Results from optimization are in Table 6 and Table 7.Table 5-Upper and Lower Bounds of Design Parameters used for Optimizationx values in µm lb ubFil_wid Outer_s Vert_cut Horz_cut4 4.257 107 117 11Inner_s Fil_thick10 170.15 0.2Table 6-Initial and Optimal Design Parametersx values in µmFil_wid Outer_s Vert_cut Horz_cut Inner_s Fil_thickInitial471110130.2Optimal Stress4.29.9111012.90.2Optimal Heat Distribution481010.7160.2Table 7-Values Obtained for the Initial and Optimal DesignValues Initial OptimalOptimal Stress (MPa) Optimal Heat Distribution (degC)92.1 88.792.2 90The stress resulting from the difference in thermal expansion in the materials was reduced from 92.1 MPa to 88.7 MPa.

This is within the yield strength of the Ultra-thin glass making this a feasible design. For even heat distribution, the initial design had standard deviation of 92.2ºC and was reduced to 90ºC. Results from the single objective optimization problems were compared as shown below Figure 7. Differences in the geometry can be seen from Table 6, but the values obtained for the solutions show little difference compared to the initial design. This may suggest that a larger design space should be used. This will enable a greater variation in results by adding or redefining the design parameters.Figure 7-Comparing Optimal Solution of Minimizing Stress with Even Heat DistributionA particular point of interest in this study was the convergence and solve time of the algorithm. The algorithm used 10 particles and 10 iterations which was determined in the study to be enough by comparison. Negligible differences in results were found for 25 iterations when compared to 10 iterations.

Negligible differences in results were found for 25 iterations when compared to 10 iterations. Even heat distribution optimization took 2 hours to solve and minimizing stress took 8 hours and 42 minutes to solve. Each iteration the overall best solution was recorded and plotted against the iteration number to show the convergence. A convergent plot shows the optimization method used arrives at a minimum of the objective function. From this we learn that this algorithm not only converges, but that it converges quickly, within 10 iterations.Figure 8-Convergence of Optimization ProblemsV. CONCLUSIONThis paper was set out to develop an optimization algorithm for designing a high temperature microheater for thermal inkjet. First, a literature review was provided on previous research on microheater design, fabrication, and operation.

First, a literature review was provided on previous research on microheater design, fabrication, and operation. A multiphysics model was developed to simulate microheater operation and evaluate microheater design, which was then validated with experimental data from literature. To optimize the microheater design, an initial geometry with set of design parameters were identified. A particle swarm optimization algorithm was then developed to search for optimal design using the experimentally validated multiphysics model to evaluate different designs. Optimal designs were found for two different objective functions, minimal stress and even heat distribution. The developed algorithm was demonstrated to be an effective and efficient tool for optimizing microheater designs. The results obtained for both objectives were shown to be very similar, which could both serve as a good microheater design for high temperature operations for thermal inkjet.VI.

The results obtained for both objectives were shown to be very similar, which could both serve as a good microheater design for high temperature operations for thermal inkjet.VI. ACKNOWLEDGEMENTSWe would like to thank my colleagues in AM3 Lab for providing support and a productive learning environment for conducting these studies and writing this paper. We gratefully acknowledge the financial support from the University of Arkansas, through the startup fund provided by the Vice Provost Office for Research and Economic Development. Any opinions, findings, and conclusions or recommendations expressed in this publication are those of the authors and do not necessarily reflect the views of the University of Arkansas.REFERENCES[1] L. Setti, C. Piana, S. Bonazzi, B. Ballarin, D. Frascaro, A. Fraleoni-Morgera and S. Giuliani, "Thermal inkjet technology for the microdeposition of biological molecules as a viable route for the realization of biosensors," Anal. Lett., vol. 37, pp.

Fraleoni-Morgera and S. Giuliani, "Thermal inkjet technology for the microdeposition of biological molecules as a viable route for the realization of biosensors," Anal. Lett., vol. 37, pp. 1559-1570, 2004.[2] W. Zhou, "Interface dynamics in inkjet deposition," 2014.[3] S. Z. Ali, F. Udrea, W. I. Milne and J. W. Gardner, "Tungsten-based SOI microhotplates for smart gas sensors," Microelectromechanical Systems, Journal Of, vol. 17, pp. 1408-1417, 2008.[4] F. Lacy, "Investigating thin films for use as temperature sensors," in Proceedings of the World Congress on Engineering & Computer Science: October 24-26, 2007, pp. 441-444.[5] L. Mele, F. Santagata, E. Iervolino, M. Mihailovic, T. Rossi, A. Tran, H. Schellevis, J. Creemer and P. Sarro, "A molybdenum MEMS microhotplate for high-temperature operation," Sensors and Actuators A: Physical, vol. 188, pp. 173-180, 2012.[6] A. Pike and J. W.

Kim, "Microheater-integrated single gas sensor array chip fabricated on flexible polyimide substrate," Sensors Actuators B: Chem., vol. 114, pp. 410-417, 2006.[18] T. Maeder, L. Sagalowicz and P. Muralt, "Stabilized platinum electrodes for ferroelectric film deposition using Ti, Ta and Zr adhesion layers," Japanese Journal of Applied Physics, vol. 37, pp. 2007-2012, 1998.[19] D. Bradley and A. Entwistle, "Determination of the emissivity, for total radiation, of small diameter Platinum-10% Rhodium wires in the temperature range 600-1450 C," British Journal of Applied Physics, vol. 12, pp. 708, 1961.[20] G. Gordillo, F. Mesa and C. Calderón, "Electrical and morphological properties of low resistivity Mo thin films prepared by magnetron sputtering," Brazilian Journal of Physics, vol. 36, pp. 982-985, 2006.[21] K. L. Chopra, "Thin film phenomena," 1969.[22] T. Demirci, D. Guney, A. Bozkurt and Y.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesign and Preliminary Evaluation of a Deployable Mobile Makerspace for Informal Additive Manufacturing Education Swapnil Sinha1, Kelsey Rieger2, Aaron D. Knochel2, and Nicholas A. Meisel1 1Made By Design Laboratory School of Engineering Design, Technology, and Professional Programs, Penn State 2School of Visual Arts, Penn StateABSTRACTAdditive Manufacturing (AM) has played an integral part in the growth of makerspaces as democratization of manufacturing continues to evolve. AM has also shown potential in enabling the successful amalgamation of art (A) with science, technology, engineering, and math (STEM) disciplines, giving new possibilities to STEAM subjects and its implementation.

This paper presents the conceptual design and development of a deployable, mobile makerspace curriculum focused on AM education for a diverse range of participant backgrounds, ages, and locations. The aim is to identify effective means of informal learning to broaden participation and increase engagement with STEAM subjects through the context of AM. The curriculum is envisioned as “material-to-form,” offering separate modules that present opportunities for self-directed learning through all the stages of design, material use, and manufacturing associated with AM. Pilot studies of the curriculum were performed to identify potential changes to improve the effectiveness of the mobile makerspace.1. INTRODUCTIONThe rapid expansion of low-cost additive manufacturing (AM, or 3D printing) systems has been one of the key catalysts in the rise of the maker movement.

INTRODUCTIONThe rapid expansion of low-cost additive manufacturing (AM, or 3D printing) systems has been one of the key catalysts in the rise of the maker movement. This movement is typified by a student desire to learn through hands-on design and fabrication experience, which maintains a primacy on sharing, connecting and do-it-yourself tinkering [1], [2]. Often, this experience occurs in increasingly common makerspaces, whether academic or commercial in nature. While makerspaces are often comprised of various forms of manufacturing (numerical milling machines, laser-cutting, welding, electronics manufacturing, etc.) many modern makerspaces also incorporate some form of AM. AM’s capacity to quickly and accurately manufacture parts of nearly unlimited complexity with minimal manual user input positions it well as a key element in learning through making. Indeed, research has shown that makerspaces are the dominant location where the majority of entry-level AM occurs [3].

Indeed, research has shown that makerspaces are the dominant location where the majority of entry-level AM occurs [3]. However, the same research discusses that, while the spectacle of the printing process often lures users to the various systems, users may maintain only a superficial or passing interest in the technology, if not encouraged to experiment with the complex relationships between the AM process and the designs that can be manufactured with it.One of the defining traits of the maker movement is that it is inherently interdisciplinary. This allows it to organically operate at the intersection of science, technology, engineering, art, and math (STEAM) disciplines. STEAM initiatives may provide an opportunity to actualize a resurgence of the fundamental importance of making to learning. This resurgence can be seen in the energy around digital media and learning [4], [5] and in the growing prominence of the maker movement.

This resurgence can be seen in the energy around digital media and learning [4], [5] and in the growing prominence of the maker movement. Increasing success of STEM students is connected to creating experiential learning opportunities. While this experiential emphasis can take many forms, utilizing interdisciplinary approaches to complex problems that draw upon STEM inquiry methods has had positive impacts on students’ self-efficacy [6]. Recently, there has been a call from federal legislatures for2801“reintegrating the two [STEM and Art disciplines] in our classrooms” [7], but the lack of substantive funding continues to marginalize initiatives in STEAM as a low priority [8]. STEAM has gained momentum as it is taken up in the popular press as a conundrum for educators [9], as a way of merging art and science education [10], and a way to “encourage holistic learning” [11].

STEAM has gained momentum as it is taken up in the popular press as a conundrum for educators [9], as a way of merging art and science education [10], and a way to “encourage holistic learning” [11]. The National Science Teachers Association (NSTA) reports incorporating Art into STEM subjects is a benefit to students and teachers in connecting concepts, exploring ideas, and increasing participation [12]. Indeed, research from the NSF funded The Art of Science Learning initiative indicates that student participants benefited from arts-based learning via greater collaboration, increased creative thinking, and longer sustained benefits in school and extracurricular participation [13].To address the opportunities inherent in utilizing makerspaces to promote informal STEAM education, this paper discusses the creation and pilot implementation of a mobile makerspace (dubbed MAKE 3D) that capitalizes on the spectacle of AM to promote design and manufacturing learning.

The system will ideally improve informal learning pathways for increasing retention and broadening participation in STEAM for students. Section 2 will discuss existing research into both the educational context to support makerspace learning as well as existing forms of AM curricular spectacle. Section 3 will then detail the design of both i) the mobile makerspace itself as well as ii) the curricular elements that scaffold the student learning experience. Section 4 will offer both quantitative and qualitative evidence of the makerspace’s impact on student engagement. Finally, Section 5 will offer conclusions and discuss avenues for future research and implementation.2. LITERATURE REVIEWRelated research relevant to this work has been divided into two sections. Section 2.1 discusses the emphasis on education to expand the field of AM and the close alignment of the maker movement to introducing informal contexts of AM to the formal context of schools.

Section 2.1 discusses the emphasis on education to expand the field of AM and the close alignment of the maker movement to introducing informal contexts of AM to the formal context of schools. Section 2.2 discusses existing development of mobile makerspaces through informal contexts such as libraries, hospitals, and museums.2.1 Educational Context to Support Informal Learning with the Maker MovementThe “2009 Roadmap for Additive Manufacturing” identified education as a critical area of research and development for advancement of the additive manufacturing field [14]. To address this critical area there have been efforts to explore AM primarily focused on university and industry training [15], but there is also interest in curriculum at the secondary level [16]. In order to grow AM learning, it is crucial to intertwine interdisciplinary curricula with simultaneous hands-on manufacturing experience.

In order to grow AM learning, it is crucial to intertwine interdisciplinary curricula with simultaneous hands-on manufacturing experience. Increasing success of STEM students is connected to creating experiential learning opportunities [17]. Makerspaces provide a set of tools within an informal context that maintains a primacy on sharing, connecting and “thinking through materials” [18].Makerspaces also are perceived to have a high level of engagement for students within an informal setting. Formal contexts, such as schools, are disconnected from students’ everyday lives and not focused on what students necessarily want to learn [19]. However, makerspaces are increasingly becoming a part of schools creating more fluid boundaries between formal and learning informal contexts for education [20]. Makerspaces provide student-centered environments that integrate technology and material play engaging more students to find value in school [21].

Makerspaces provide student-centered environments that integrate technology and material play engaging more students to find value in school [21]. When students create with technology they “become more engaged, spend more time2802investigating and/or constructing and take ownership for and build confidence in their abilities to learn and understand” [22]. Part of what creates engagement in the makerspace context is the importance of creativity and play with materials, often expressed through the concept of tinkering. Tinkering is an open-ended process of inquiry that involves improvisation and creativity [23].2.2 Existing Forms of AM Curricular Spectacle and Mobile MakerspacesMakerspaces combine informal learning utilizing AM as a central fabrication technology, interdisciplinary approaches in STEAM, and investigations into materials, but are often situated within a particular space, which limits their impact.

MAKE 3D adds an additional aspect through the mobility of our platform extending what we call the curricular spectacle of our approach. Curricular spectacle refers to educational efforts that involve highly visible or novel introductions to content or technologies that engage learners immediately while possibly leading to deeper understanding. Attempts in recent years to capture this spectacle-driven fascination with AM technology to informally guide users to a deeper understanding of manufacturing and design include MakerBot Innovation Centers and 3D printing vending machines [24]. Beyond these static forms of informal AM making, libraries, universities, and K-12 schools have also been experimenting with the use of mobile makerspaces to create a sense of spectacle that can be easily transported from location to location (see, for example, [25]–[28]).

While these mobile makerspaces have been shown to excite local communities and create a sense of wonder regarding AM technology, there have been no studies to assess whether or not such mobile makerspaces are capable of taking the user’s initial spectacle-driven fascination and sustain it into learning and engagement with the STEAM disciplines that rely on the concept of making.Research on informal learning and makerspaces has demonstrated its import for science education [29], [30], and much of this work in informal spaces has focused on libraries and museums [31]–[33]. Library programs such as the Arrowhead Library System mobile makerspace, started in 2014, to develop a model of shared equipment relevant for making among seven small to mid-size libraries in southern Wisconsin (see http://als.lib.wi.us/Makerspace/). There has also been mobile making in hospitals to explore its therapeutic capacity such as Gokul Krishnan’s mobile makerspace at Monroe Carell Jr.

There has also been mobile making in hospitals to explore its therapeutic capacity such as Gokul Krishnan’s mobile makerspace at Monroe Carell Jr. Children's Hospital at Vanderbilt in Nashville, Tennessee [34] which (see http://www.makertherapy.com/). All of these programs are engaging with issues of maker education in ways that is mobile, open, and accessible to all. Our research is intended to add to this literature with the unique capacities of mobile makerspaces with adaptable curricular kits that engage learners in design and material education using AM to extend this high-visibility approach and better understand the effects of heightened awareness, hands-on learning, and broader participation in making as an experiential approach to STEAM disciplines.helpstoaddresstheneedsofchronicallyillpatients3. DESIGN APPROACH TO THE MOBILE MAKERSPACEThe essence of a mobile makerspace lies in its nature of being accessible, both physically and intellectually [35].

DESIGN APPROACH TO THE MOBILE MAKERSPACEThe essence of a mobile makerspace lies in its nature of being accessible, both physically and intellectually [35]. Just like the mobile libraries or the bookmobiles which were introduced to provide library services everywhere [36], the mobile aspect of a makerspace ensures the physical accessibility of a learning environment. Additionally, a self-driven and experiential learning offered by these makerspaces is expected to improve participation in STEAM [37]. Subsequently, the MAKE 3D truck is intended to be built so as to offer effective learning and retention of knowledge offered around AM. To meet this objective, the design of the trailer dedicated as the2803mobile for this deployable makerspace, and its operation is elaborated in the section 3.1.

To meet this objective, the design of the trailer dedicated as the2803mobile for this deployable makerspace, and its operation is elaborated in the section 3.1. The development of the curriculum for the same is explained in the section 3.2.3.1 Overview of the MAKE 3D Space and Educational ApproachThe entirety of the deployable MAKE 3D makerspace is contained within a single mobile platform. The platform chosen for this implementation is a closed-wall trailer (see Figure 1) with interior space of 132” x 72” x 78”. The different equipment, consumables, and support systems necessary to operate the makerspace were custom designed and selected in order to best make use of the mobile environment. When the MAKE 3D trailer arrives at the desired location, all contents can be removed and deployed in a flexible configuration.

When the MAKE 3D trailer arrives at the desired location, all contents can be removed and deployed in a flexible configuration. This allows it the unique ability to have the learning space cater to the needs of the student, rather than requiring the student to conform to the limitations of the learning space.(a)(b)Figure 1. MAKE 3D trailer with (a) side view, and (b) back view.The curriculum for MAKE 3D was developed with the intention of introducing AM in a makerspace platform to test and advance its role in improving the effectiveness of informal learning pathways. Via a comprehensive spectrum of rudimentary concepts referred to in this project as the “Material to Form” model, the curriculum for MAKE 3D centers around AM as a fabrication method in making from an interdisciplinary and layman understanding. As the term “material to form” suggests, the curriculum for MAKE 3D was originally conceived of as a comprehensive project with a clear start-to-finish trajectory.

As the term “material to form” suggests, the curriculum for MAKE 3D was originally conceived of as a comprehensive project with a clear start-to-finish trajectory. However, given the nature of the makerspace as an informal learning resource for voluntary learners and the project’s interdisciplinary platform, it became apparent that the materials and curriculum would need to be accessible from multiple entry points. As such, each learning module required flexibility to address varying expertise, learning styles, time constraints, and levels of interest. This is done through various demonstrative pieces that invites participants to try on hands-on activities that emphasizes design and inquiry-based methodologies like Design Thinking [38] as integral components of making. Throughout the development of the MAKE 3D project, the Material to Form curriculum has gone through multiple iterations informed by ongoing pilot studies.

Throughout the development of the MAKE 3D project, the Material to Form curriculum has gone through multiple iterations informed by ongoing pilot studies. It was determined early on that these elements could be grouped into six categories of modules in the curriculum, based on the application and equipment required. The modules included Designing Form, Modeling Form, Capturing Form, Workflow in AM, Extrusion, and Filament Variety. These modules will be discussed in more detail in Section 4.2804Each module was initially developed as self-contained units from start to finish, in order to keep in facilitate a more non-sequential and informal flow of participants through modules. Additionally, each module was developed with its own set of equipment, a handout, and a set of slides. In the early stages of development, equipment for each module includes all the possible hardware, software, and the exemplary relevant 3D printed pieces to be included in a makerspace station.

In the early stages of development, equipment for each module includes all the possible hardware, software, and the exemplary relevant 3D printed pieces to be included in a makerspace station. As the build of the makerspace progresses curricular modules were expressed in makerspace stations with some overlap. For example, the Capturing Form and Workflow in AM curricular modules are both facilitated in the makerspace computer station. Handouts include sets of directions for activities for each module in step-by-step screen shots and short sentences, ending with reflection questions prompting participants to think about what they learned through the activities (see Figure 2). Finally, slide shows were created for each module as an additional resource for the instructor to help explain concepts, if necessary.

Finally, slide shows were created for each module as an additional resource for the instructor to help explain concepts, if necessary. However, in order to maintain the informal nature of the approach, slides are maintained on individual tablets, rather than projected on a large screen for the entire group.Figure 2. Examples of handouts in the Work Flow, Capturing Form, and Modeling Form modulesA gimbal design (see Figure 3) was created to serve as a base example for the activities in all six modules. This geometry serves as a tool to provide scaffolding for student learning throughout the makerspace. The gimbal is defined as a pivoted support that allows the rotation of an object around a single axis, and thus requires participants to consider its aesthetic and functional aspect. [39]. This assisted in keeping the modules tied to each other, creating a flow between them.

[39]. This assisted in keeping the modules tied to each other, creating a flow between them. For example, geometry serves as a base of Design Form module by prompting students to think and create other forms of gimbals, it also serves as a base for Modeling Form to modify it in the digital platform, etc. A common article, like the gimbal, is expected to provide participants with a motive to proceed to other modules, by introducing the possibilities of variations that can be performed on it. The gimbal was chosen because of the spectacle it offered with its function. Its fabrication with 3D printing as an assembled part in one build adds an additional wonder to the possibilities of AM, which may invite curiosity and consequently self-motivated involvement in the modules.2805Figure 3.

Gimbal design created as a common artifact for continuity between curricular modules.3.2 Details of Individual Educational ModulesThe aim for MAKE 3D is to provide hands-on learning experience in AM to a virtually unexposed group of students. To successfully create a platform providing basic understanding of AM, the curriculum was built around the equipment providing demonstrative, and free form interaction. The curriculum covers design thinking, digital designing, digital capturing or scanning, work flow in AM (from conversion of files into .STL format to slicing it for desired part quality), extrusion (with 3D pens and clay extruders), and filament variety available.

These elements were disseminated into different modules according to the hardware and software utilized, and are elaborated below.Designing FormAs the only learning module not directly tied to a particular element of AM machinery, designing form is a creative prototyping activity based on the process of Design Thinking [38]. The conceptual foundation for this module is based on the understanding that design is an iterative process, which can be strategically performed using the five design thinking stages. These stages are: empathize, define, ideate, prototype, and test. From the activities stipulated in the Designing Form curriculum, MAKE 3D participants learn to use this creative and critical process to effectively design a personalized product, such as the gimbal coin discussed earlier. To support participant engagement with this process, nontechnical materials such as cardboard, scissors, and tape are used for the prototyping process.

To support participant engagement with this process, nontechnical materials such as cardboard, scissors, and tape are used for the prototyping process. Participants begin by sketching various versions of their design (see Figure 4), before creating cardboard rotating prototypes in the form of their initial design. The Designing Form module was set up as an individual station that includes all the necessary prototyping supplies and tools.Figure 4. A participant engaged with Designing Form handout, creating gimbal inspired designs and participants engaged in prototyping.2806Modeling FormThe digital modeling component of the curriculum integrates opportunities for experimentation with available programs such as Fusion 360 (https://www.autodesk.com/), Tinkercad (https://www.tinkercad.com/), and SketchUp (http://www.sketchup.com/).

Each of these platforms were considered to determine the most user-friendly and accessible program available for MAKE 3D participants to engage in designing new forms. Tinkercad was ultimately chosen for the makerspace due to the fact that it is free to the public and promotes a simplified, straightforward set of commands. This module provides participants with the opportunity to learn by doing, and may or may not be based on their previously prototyped gimbal design depending on preference and interest (see Figure 5). The handout guides them to create a personalized keychain by importing an existing gimbal design and modifying it. The directions aim to create an understanding of spatial orientation in digital design, and how to manipulate it to create desired designs. The Modeling Form module was setup along with the Work Flow module as a grouped station, since both requires computer for hands-on activities. Six computers were setup on a table of dimension 72” x 32”.Figure 5.

Six computers were setup on a table of dimension 72” x 32”.Figure 5. Use of computer station to learn digital design in Modeling Form module.Capturing FormSimilar to Modeling Form, the scanning equipment provided in the Capturing Form Module is meant to offer participants the opportunity to engage freely with available technology in order to better grasp its capabilities. This engagement occurs via Occipital Structure Sensors on iPad, SCANN3D application for android phones, and Microsoft Kinect Scanner programs that are demonstrated for the students prior to use. This hand-on engagement allows participants to directly observe how 3D scanners are able to capture surface data to be converted into a printable .STL file (see Figure 6), and to educate participants in the area of mass customization as one of the driving forces in design for AM. The Capturing Form module was set up as an individual station, with the Structure Sensor and the Kinect on a small table.

The Capturing Form module was set up as an individual station, with the Structure Sensor and the Kinect on a small table. A rotating turntable of diameter 32” was constructed as a scanning rig for a hands-on good quality 3D scanning with Kinect. One of the computers in the modeling form and workflow station was dedicated for the Kinect scanning.2807(a)(b)Figure 6. Scanning station with equipment in use (a) Structure Scanner being used by the participants to scan a hand, (b) Experts demonstrating use of the turntable to obtain a good quality 3D scan with Kinect.Work Flow in AMThe Work Flow module focuses on the print preparation procedure from importing the digital design in STL file formats to exporting the 3D print-ready file. This module demonstrates the different parameters in the slicing software (Cura) that can be varied to create the desired output.

This module demonstrates the different parameters in the slicing software (Cura) that can be varied to create the desired output. It prompts participants to think about considerations like support material, infill densities, layer thicknesses and their impact on part quality and total time for print. The layer animation of the previewed file demonstrates the layer-by-layer addition of material in AM, explaining the capability of fabricating complex part. The workflow module also incorporates demonstrative 3D printed objects to show the need for support material, layer thicknesses, infill densities and patterns, and various slices of an object printed as a stacking puzzle game.

This module introduces participants with the restriction as well as opportunities of AM, an important aspect to be considered while designing for AM.Extrusion and Filament VarietyMaterial extrusion systems in the form of 3D pens, handheld clay extrusion tools, 3D printers, and filament extruders are leveraged alongside 3D printers using both clay and plastic filament. The inclusion of a filament extruder also provides an initial introduction to the concept of raw material processing within AM. This simple system demonstrates the ways in which plastic pellets are heated, extruded, and wound around a spool to form the processed material for material extrusion AM (see Figure 7). These materials are used in collaboration with an information handout and worksheet to allow participants to directly observe the manufacturing phase for their designed products by relating the complexities of their chosen design to the capabilities of the manufacturing system.

Furthermore, these materials help participants conceptualize the physical capabilities of the 3D printer and the potential for expanding the X, Y, and Z planes of a two- dimensional object. Along with the learning goals of the Workflow Module, these curricular materials engage participants in the process of producing a three-dimensional object using variety of materials with AM. The hands-on involvement helps in the most effective demonstration of AM extrusion process behavior with different materials like metal, wood, flexible filaments, and conductive materials. Extrusion module and the Filament Variety module was set up together on a table along with the 3D extrusion Pens and the clay extruders. Various filaments were available for use with 3D extrusion Pens. Multiple material extrusion printers were setup together along with a demonstrative clay printer.2808Figure 7.

Various filaments were available for use with 3D extrusion Pens. Multiple material extrusion printers were setup together along with a demonstrative clay printer.2808Figure 7. Participants creating 3D objects with 3D Pens at extrusion station, playing with different filament variety.Demonstration GalleryFinally, while not a formal curricular module, a Demonstration Gallery (see Figure 8) was conceived to add a collection of 3D printed items, which acted as the demonstration for each module. This gallery acts as a curated collection of engineering design, science, mathematics, and art objects in a range of printed materials that are meant to illustrate the wide variety of potential part applications, surfaces, densities, and AM capabilities. Objects included in this gallery are considered case studies, acting as examples of the themes found in each of the established learning modules.

Objects included in this gallery are considered case studies, acting as examples of the themes found in each of the established learning modules. Due to the open source nature of much of the maker culture, all of the works featured in the gallery include corresponding .STL files and resources for reproduction that are accessible to the participants and the general public. While all these modules were developed around material extrusion process in AM, owing to its simplicity, accessibility, and robustness for an educational setting, the other major process types in AM are demonstrated in form of example prints showcased in the Demonstration Gallery.Figure 8. Examples demonstrated in the Display Gallery4. PRELIMINARY USAGE CASE STUDYTo test the applicability of the curriculum and the equipment developed for the MAKE 3D, a pilot study was performed on four undergraduate class sections, each comprised of approximately 24 students.

The aim was to observe the functioning of each module as an individual station, in terms of interaction by the participants, and the involvement of instructors. Additionally, the effectiveness of the curriculum in boosting participant’s comprehension of makerspaces, design thinking, and AM technology was observed.28094.1 Educational ContextFor the pilot studies, a room was setup as a preliminary mock-up of the deployed form of MAKE 3D with all the six modules in the curriculum setup as stations (see Figure 9 and Figure 10). The stations were grouped as explained in Section 3.1, according to the equipment requirements and named as Designing, Computers, Extrusion, Scanning, Printing, and the Gallery for data collection purposes. While Designing, and Scanning were individual stations for the Designing Form module and the Capturing Form modules, respectively, the Computer station was a combined setup for the Modeling Form and Work Flow modules.

Similarly, the Extrusion station was a combined setup for Extrusion module as well as the Filament Variety module. The printing station consisted of five material extrusion printers (Monoprice Mini) for free use by the participants, as well as a clay printer for demonstration. The Gallery station was named for the gallery display. Each station included packets consisting of the collected handouts for all modules, with posters at each station also included to convey the same information in a larger format. Experts in AM informally moved between stations to guide or demonstrate concepts or equipment to the participants if required.Figure 9. Conceptual arrangement of the deployed MAKE 3D set up for pilot study, with connections between physical stations (rectangles) and corresponding modules (ovals).Figure 10.

Conceptual arrangement of the deployed MAKE 3D set up for pilot study, with connections between physical stations (rectangles) and corresponding modules (ovals).Figure 10. A portion of deployed MAKE 3D makerspace stations present at the pilot studies.2810Before exposure to the deployed MAKE 3D space, the participants were prompted to self-rate their agreement on familiarity with makerspaces, 3D printing, and design thinking, etc. on a Likert scale. The participants were given a short premise of MAKE 3D as a makerspace, before the exposure. Two raters individually and simultaneously completed observation sheets with a narrative description of what they perceived was the participant’s response and reaction to each station every 10 minutes from the beginning to the end of the session. Additional data for number of participants and their nature of interaction with the module was also collected for each station.

Additional data for number of participants and their nature of interaction with the module was also collected for each station. Each session lasted for approximately one and a half hours.4.2 ResultsThe age of the 95 participants ranged from 17 to 20, (15 females and 80 males). These participants were all students participating in an introductory engineering design course targeted at first-year engineering students. The participants were asked to finish a 20-question survey on a Likert scale (see Figure 11) right before their exposure to makerspace. The same survey was taken right after the exposure. A sign test with continuity correction was used to compare differences in the Likert scale for each question, pre- and post-exposure, using IBM SPSS Statistics 24.Figure 11.

A sign test with continuity correction was used to compare differences in the Likert scale for each question, pre- and post-exposure, using IBM SPSS Statistics 24.Figure 11. Likert Scale as provided to the participants for each question, pre and post-exposure.The mean values of the participant ratings improved for 19 of the 20 survey questions post exposure, compared to their pre exposure data. To find the significance in improvement, statistical analysis was performed within subjects. On statistically analyzing the pre-exposure and post- exposure student surveys for the same 20 questions, a statistically significant increase was found for 10 of the 20 questions (questions with a significant increase are shown in Figure 12). Three questions in particular saw drastic improvement. First, it was found that students rated themselves statistically significantly more familiar with concepts of a makerspace post exposure, (p<0.0005, increase from “Disagree” to “Agree”).

First, it was found that students rated themselves statistically significantly more familiar with concepts of a makerspace post exposure, (p<0.0005, increase from “Disagree” to “Agree”). This metric gives us a self-reported engagement with MAKE 3D as a complete system, where a significant improvement in rating indicates a good level of engagement in the space. Students also rated higher on use of 3D printing, post-exposure, compared to pre-exposure (p<0.0005, increase from “Neutral” to “Agree”). The jump from neutral to agreement means that the makerspace is properly executing one of its main goals: offering students a chance to go hands-on with 3D printing technologies in order to improve their engagement with this technology. Additionally, a significant increase was reported in the students’ understanding of 3D scanning technology (p<0.0005, increase from “Disagree” to “Agree”).

Additionally, a significant increase was reported in the students’ understanding of 3D scanning technology (p<0.0005, increase from “Disagree” to “Agree”). Similar to the use of printers, this jump from disagreement to agreement indicates that the curriculum was able to showcase the use of 3D scanning in a way that is relevant to the participants.2811Figure 12. Average Likert scale rating for survey questions with a statistically significant change between pre- and post-exposure. Questions with the most drastic changes are highlighted on the vertical axis in a box.Two observers simultaneously recorded the descriptive narration of what they observed in each station in every 10 minutes. They recorded the number of students at the station, and overall engagement in the scale of 1 to 5, and type of interaction with the modules for each station. Reliability for the engagement level rating scale from two Raters at each station was found by calculating Chronbach’s alpha.

Reliability for the engagement level rating scale from two Raters at each station was found by calculating Chronbach’s alpha. The test was performed with IBM SPSS Statistics 24, on data points with common time and common number of participants. The test revealed high reliability of the scale with Chronbach’s alpha equating to 0.96. On observing the average for the ratings obtained from the two raters on engagement, it was found that the computer station with the Modeling Form module was the most engaged station, followed by the extrusion station with 3D pens and clay extruders with variety filament, further followed by Printing, Design, Scanning, and the Gallery stations respectively (see Figure 13). Additionally, it was observed that extrusion had the highest count of participants every 10 minutes followed by Computer, Printing, Design, Scanning, and Gallery respectively.Figure 13.

Additionally, it was observed that extrusion had the highest count of participants every 10 minutes followed by Computer, Printing, Design, Scanning, and Gallery respectively.Figure 13. Average engagement rating by the raters for each station, along with the average count of participants every 10 minutes in each station.2812Observation notes support the engagement ratings in clearly highlighting the popularity of the extrusion table often citing freeform design with the 3D pens and even extruding designs onto objects such as smartphone cases and older discarded prints that were made available. Peer interaction and numbers of participants remained high throughout the session suggesting the success of the informal environment of tinkering, but uncertain as to how the module connected with or motivated students to move to other stations to work on other curricular modules.

Observation notes also confirmed that participation at the computer station appeared to be longer term for certain users as it was noted that participants would stay on the computer station designing 3D forms in Tinkercad for long periods of time which relates to both levels of engagement but also the time intensive practice of 3D modeling. It is unclear as to how the computer station related to other modules, because there were observations indicating that participants were progressing from designing form to utilizing the 3D printing workflow to prepare their prints. However, improper build profiles loaded on the 3D printers would sometimes hamper the production of prints and made it difficult to track this production cycle.

However, improper build profiles loaded on the 3D printers would sometimes hamper the production of prints and made it difficult to track this production cycle. Observation notes indicating low participation in the Design and Gallery stations highlighted the inadequate curricular supports provided in the paper handouts and lack of instruction for each module to provide entry points to making at the various makerspace stations. Revisions to the makerspace and curriculum will take these considerations into account as we make changes for subsequent deployments.5. CONCLUSIONS AND FUTURE WORKThe MAKE 3D mobile makerspace is conceived as a platform for individuals to gain hands- on experience with AM technology.

CONCLUSIONS AND FUTURE WORKThe MAKE 3D mobile makerspace is conceived as a platform for individuals to gain hands- on experience with AM technology. It is built based on a “material to form” curriculum around basic equipment such as 3D extrusion pens, clay extruders, 3D scanning devices, software to create or manipulate digital designs, prepare 3D print file while gaining an understanding of parameters that can impact their print quality and print time. These tools are tied with a design thinking process that introduces a novice to a methodological and effective approach to design. The hope is to showcase the applications of this technology as well as to familiarize participants with techniques available to use it for their projects or hobbies. A pilot study was run to observe the effectiveness of the design of the curriculum in communicating the accessibility and simplicity of AM technology for use.

A pilot study was run to observe the effectiveness of the design of the curriculum in communicating the accessibility and simplicity of AM technology for use. Analysis on self-reported survey ratings showed that participants were able to successfully engage with the MAKE 3D as a makerspace, and were able to better understand how to use both 3D printing and 3D scanning tools at their disposal.Observations on participant engagement level and their involvement at the stations showed us that some stations grabbed more attention compared to others. To address these variations and further expand upon participants’ hands-on experiences, it was determined that a series of revisions were needed to enable more robust participation across the curricular modules and corresponding makerspace stations. First, curricular materials will be extended to include digital copies at computer stations make content more readily accessible.

First, curricular materials will be extended to include digital copies at computer stations make content more readily accessible. Second, the arrangement of the stations will be deployed in variations that will couple under used and over used stations that test proximity effects. For example, the extrusion station will be deployed in the vicinity of the gallery to find out how the interaction of the modules may effect participation. Last, the research team will develop a schedule of mini demonstrations that will support participants in more technical endeavors while trying to maintain the overall informal context of the makerspace. The next steps for MAKE 3D is to tour four campuses of the Penn State University system along with other possible informal2813contexts such as the Maker Week at Schlow Centre Region Library in State College, PA.

Through the revision process, the research team will remain focused on the fundamental underlying processes and techniques of AM while also maintaining learners’ engagement through the informal opportunities of the makerspaces using curricular spectacle.6. REFERENCES [1]S. Libow Martinez and G. S. Stager, “Invent to learn: Making, tinkering, and engineering in the classroom,” p. 237, 2013. [2] J. Walter Herrmann and C. Büching, Fablab of machine, makers and inventors. 2013. [3] C. Bosqué, “What are you printing? Ambivalent emancipation by 3D printing,” RapidPrototyp. J., vol. 21, no. 5, pp. 572–581, 2015.[4] Y. B. Kafai, K. A. Peppler, and R. N. Chapman, The Computer Clubhouse: Constructionism and Creativity in Youth Communities. Technology, Education--Connections. ERIC, 2009. [5] Y. Kafai and M. Resnick, Constructionism in practice: Design, Thinking, and Learning ina Digital World. Routledge, 1996.[6] A.

Trends, vol. 9, no. 3, pp. 117–119,1961.[37] L. Martin, “The Promise of the Maker Movement for Education,” J. Pre-College Eng. Educ.Res., vol. 5, no. 1, 2015.[38] P. G. Rowe, Design thinking. MIT press, 1991. [39] T. Atkins and M. Escudier, Dictionary of mechanical engineering, no. 1. Oxford UniversityPress, 2014.7. ACKNOWLEDGEMENTSWe gratefully acknowledge the support of the National Science Foundation Grant No. 1623494. Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the views of the National Science Foundation. We would also like to thank Tom Lauerman for his significant efforts in the design and implementation of the MAKE 3D space.2815.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDESIGN AND PROCESS CONSIDERATIONS FOR EFFECTIVE ADDITIVE MANUFACTURING OF HEAT EXCHANGERSEvan Handler1, Amanda Sterling2, Jonathan Pegues2, Huseyin Ozdes2, Mohammad Masoomi2, Nima Shamsaei2, Scott M. Thompson2,†1Naval Surface Warfare Center, Carderock Division, Potomac, MD 20817 USA2Laboratory for Fatigue and Additive Manufacturing Excellence, Department of Mechanical Engineering, Auburn University, AL 36849 USA† Corresponding author: Email: smthompson@auburn.edu Phone: (334) 844 4867AbstractThis paper provides some insights into using powder bed fusion (PBF) techniques for additively manufacturing heat transfer equipment (HTE), such as heat pipes and heat sinks. Background information is provided on the operating principles of PBF and the subsequent features of parts fabricated via PBF.

Background information is provided on the operating principles of PBF and the subsequent features of parts fabricated via PBF. Examples of heat transfer equipment produced using PBF are discussed. Some benefits and challenges associated in using PBF for generating effective heat transfer equipment are summarized.IntroductionIn recent years, there has been interest in using powder bed fusion (PBF) for the additive manufacture (AM) of heat transfer equipment (HTE). PBF methods provide several advantages over traditional manufacturing methods in this area. The most commonly claimed advantage is its ability to create objects with geometries once impossible to create with traditional manufacturing processes.

The most commonly claimed advantage is its ability to create objects with geometries once impossible to create with traditional manufacturing processes. PBF enables the ability to create complex internal structures in a single manufacturing step, whereas most heat exchangers that are designed with internal structures must be assembled using various joining methods that can cause thermal resistance and decrease heat transfer rates.While there are many different types of HTE, they all generally encounter similar challenges in the manufacturing process. One example is the issue of thermal resistance between adjoined surfaces, which typically is caused by small gaps between the surfaces, though the issue also is present when two surfaces are joined using traditional joining methods.

Furthermore, since research related to the design of THE is often done computationally to create highly- optimized designs [1], computer-generated designs sometimes can be impractical or even impossible to manufacture. Furthermore, as miniaturization becomes increasingly prevalent, careful consideration must be given toward manufacturing methods required for parts. Thus, for every advance made in thermal management devices, its benefit must be weighed against the cost of manufacturing and implementing new components.2632The AM of a part is accomplished through guidance of a CAD file and by adding material continuously, in a layer by layer fashion, until a finished product is produced in a usable or near usable state. Among the different AM techniques, PBF has become the most commonly researched and used technique. PBF works by using a moving heat source that is supplied either by a laser or an electron beam.

Among the different AM techniques, PBF has become the most commonly researched and used technique. PBF works by using a moving heat source that is supplied either by a laser or an electron beam. The heat source is controlled by computer numerical control (CNC) to melt specific areas in a bed of metallic powder to form a single layer. Then, the build stage is lowered, and another layer of powder is coated on top of the previous layer; this process is repeated until the final part is produced. Laser-based PBF is sometimes referred to as laser- powder bed fusion (L-PBF) selective laser melting (SLM) or direct metal laser sintering (DMLS).PBF has several advantages over other metal-based AM methods, including smaller layers, finer details, and the ability to create internal structures without supports over the previous layers of unmelted powder.

PBF has been shown to be effective in producing parts with a wide variety of metal alloys, including many different titanium-, steel-, aluminum-, nickel- and commercially pure copper-based alloys [6,7]. In industry, PBF has been used successfully to create various parts, such as the fuel injector for a new General Electric (GE) jet engine [4], a copper combustion chamber for NASA [5], and even biocompatible tissue scaffolds [6].Surface RoughnessPBF methods are known to produce rough surfaces on manufactured parts. Under many circumstances, these surfaces are either machined or sandblasted to make the surfaces smoother, but some researchers have chosen to study the roughness of the surfaces with the goal of finding ways to reduce or prevent it. While rough surfaces are undesirable in structural and moving parts that undergo fatigue, they can be both a benefit and a hindrance in heat exchangers.

While rough surfaces are undesirable in structural and moving parts that undergo fatigue, they can be both a benefit and a hindrance in heat exchangers. For example, roughness has been shown to increase heat transfer by causing more local turbulence in the working fluids and providing more surface area, but it also has been shown to increase the pressure drop across surfaces.Rough surfaces are a direct result of the PBF process, and they can be caused by various mechanisms. The most visible surface roughness is caused by the successive, partially- overlapping passes of the heat source, which causes successive, visible grooves between the path of the heat source and the successive layers. This is due to the influence of surface tension, which causes the top of the molten pool to take a spherical shape.

This is due to the influence of surface tension, which causes the top of the molten pool to take a spherical shape. Furthermore, early in the process, this effect of surface tension can cause some molten particles to “splash” (or spatter) and then solidify as much smaller particles if they land on a previously-melted section. It is common for the surfaces of particles just outside the melt pool to partially melt and become sintered to the part as the heat source moves away [7]. An example of the surface roughness inherent to PBF surfaces/channels is shown in Fig. 1 which shows a channel structure next to a milled surface. The surface was milled to visualize mini-channel surface roughness within a PBF/Ti-6Al-4V flat oscillating heat pipe [8].Most studies have indicated that surface roughness is dominated, in large part, by the process parameters that exist when the basic samples are produced.

Hatch spacing directly influences the distance between grooves, and lower hatch spacing decreases this distance [16- 18]; the thickness of the layers causes a similar effect in the z direction, although this does not occur in DMLS [9]. The direction in which the surface roughness builds is another important2633factor because vertical and horizontal surfaces will have drastically different surface roughnesses [17,19]. Scan velocity and the power of the heat source have been shown to have some influence on surface roughness because they influence both the heating rate and subsequent cooling rate, but these influences are not always significant [17, 18, 20]. The size of the particles that make up the powder and their size distribution have significant effects on surface roughness; smaller particle sizes and smaller size distributions in the powder decrease surface roughness [16, 21].

The extent to which all these parameters effect surface roughness varies based on the method and the material used. Some parameter optimization studies have attempted to determine sets of parameters that result in the lowest surface roughness. Calignano et al. [12] investigated the effects of different SLS process parameters on the surface roughness of AlSi10Mg, and they used the Taguchi method to determine an ideal set of parameters for the best surface roughness. Optimal parameters were chosen based on this analysis, and the samples that were produced using these parameters were found to have an average surface roughness of Ra = 15.68 μm.Figure 1. SEM image of a PBF/Ti-6Al-4V mini-channel adjacent to a machined surface.The geometrical orientation of the part during its AM has a substantial influence on surface roughness, especially when complex geometries are involved. Pakkanen et al.

Pakkanen et al. [14] studied this phenomenon using SLM to produce variations of tubing 10 mm thick out of AlSi10Mg and Ti-6Al-4V. The AlSi10Mg tubes were produced at various angles between the horizontal and vertical orientations, while the Ti-6Al-4V tubes were produced only at an angle of 45º. The tubing samples were cut in half, and surface roughness measurements were taken from the top and bottom of the internal channels for comparison. All cases were found to have Ra values of 20 μm with an error of ±5 μm.There are other phenomena and circumstances that affect surface roughness in different ways. For example, Jamshidinia et al. studied the effect of heat buildup on surface roughness in EBM by producing sets of two thin walls made of Ti-6Al-4V with different spacing between them [15].

For example, Jamshidinia et al. studied the effect of heat buildup on surface roughness in EBM by producing sets of two thin walls made of Ti-6Al-4V with different spacing between them [15]. They found that lower heat accumulation resulted in lower roughness values, and they also found that there was an approximate logarithmic relationship between the spacing distance and surface roughness. Wang et al. used SLM to produce curved overhang structures out of SS 316L with different laser power settings to study the surface roughness of both the top and bottom surfaces as the incline angle increased [16]. They found that surface roughness began to increase sharply when the incline angle exceeded 45º for both surfaces.

They found that surface roughness began to increase sharply when the incline angle exceeded 45º for both surfaces. Distinct changes in the2634surfaces were observed on all samples as they transitioned from fully dense and successful to build failure.Porous StructuresPorous materials often are used in heat pipes as wicking structures, often implemented through metal foams or sintering processes. However, many ‘traditional’ porous foams involve a high degree of randomness, and their effectiveness cannot always be considered optimal. In comparison, PBF methods provide an opportunity to manufacture porous structures with pre- optimized designs for more efficient HTE. The work of Bodla et al. numerically showed that well-designed artificial porosity can produce significant increases in effective heat transfer over other traditionally-produced materials [17]. The production and use of cellular lattice structures have been investigated.

The production and use of cellular lattice structures have been investigated. These are beneficial in increasing heat transfer due to their high surface area and continuous nature. To date, many different types of cells have been additively manufactured successfully, and much work has been done to investigate the manufacturability of different volume fractions and the changes in the strut and pore sizes [31 - 33]. Some shapes have been tested and confirmed to have higher effective heat transfer coefficients than those in otherwise porous materials [20]. The design of these structures currently is limited by the minimum producible feature size [20].Heat SinksWong et al. used SLM to produce heat sink arrays of Al 6061 and some SS 316L with different geometries to compare their associated heat transfer and pressure drop during cross- flow of air through the arrays [34 - 36].

The geometrical shapes that were produced included cylindrical, rectangular, diamond-shaped, and elliptical pins, as well as an alternating “V” array and a lattice array – with an example of the lattice array shown in Fig. 2. In all of the studies, the heat sinks were manufactured on a 50 x 100-mm substrate of the same material used in the production of the fins. Testing was performed using a 16-mm copper block heated by two 6-mm diameter electrical cartridge heaters bolted to the bottom surface of the heat sinks. Ambient air was forced through piping at varying velocities, and the lengths of the pipes were designed for fully-developed conditions both before the heat sink and before the final measurements were taken. The heated section was fully insulated. For comparison, pressure and temperature were measured before and after the heat sink. In all cases, surface roughness was estimated to be between 15-25 μm.2635Figure 2. Lattice-structured heat sink fabricated via L-PBF [34].

In all cases, surface roughness was estimated to be between 15-25 μm.2635Figure 2. Lattice-structured heat sink fabricated via L-PBF [34]. Used with permission.When SS 316L heat sinks were tested, their performances were only about half those of the Al 6061 heat sinks due to the poor thermal conductivity of SS 316L [23]. When the different geometries were compared, the cylindrical pins and the “V-shaped” array performed roughly the same at equal Reynolds numbers, and the diamond-shaped pins produced significantly higher heat transfer at much lower Reynolds numbers, while the elliptical pins produced the lowest heat transfer. However, the cylindrical pins resulted in a nearly constant pressure drop, and the diamond-shaped pins initially created higher pressure drops that decreased as Reynolds number increased.

However, the cylindrical pins resulted in a nearly constant pressure drop, and the diamond-shaped pins initially created higher pressure drops that decreased as Reynolds number increased. The “V-shaped” array produced significantly lower pressure drops that decreased as Reynolds number increased until eventually leveling out, and the elliptical pins fell roughly in the middle [36, 37]. These results changed when rectangular pins and lattice structures, modified cylindrical pin dimensions, and modified elliptical pin dimensions were tested. With better dimensions, the rectangular and elliptical pins were found to have similar heat transfer when tested at the same flow rates, but the cylindrical pins still were inferior. However, the lattice structure had significantly lower heat transfer even though it had a larger surface area.

However, the lattice structure had significantly lower heat transfer even though it had a larger surface area. This was due in part to the fact that the lattice was designed to have channels that allow air flow with little interaction with the lattice structure, and this reduces heat transfer. When the pressure drops were compared, the elliptical pins produced the lowest pressure drops, the cylindrical pins produced the highest, and the pressure drops for the rectangular pins and lattice structure were about the same.Dede et al. used 3D topology optimization methods to generate a geometry file for a pin- based heat sink, which was manufactured by SLM using AlSi12 [1]. Error! Reference source not found. shows the CAD geometry of this optimized heat sink. Then, the printed heat sink was tested against heat sinks made out of Al 7075 and oxygen free Cu and produced using traditional manufacturing techniques.

Then, the printed heat sink was tested against heat sinks made out of Al 7075 and oxygen free Cu and produced using traditional manufacturing techniques. These heat sinks were tested using air jet infringement, and the pressure drops and air temperatures were recorded. The heat sinks were mounted to a heater, which was backed by PEEK insulation.During the experiments, the air flow rate was set and the heater was ramped up until there was a temperature difference of 15 K between the heater and the average temperature of the surface-mounted thermocouples. The system was allowed 30 min to reach steady state before the temperatures and pressure drops were recorded. The temperature and pressure data were used to2636calculate the thermal resistance. Two different cases were studied, and the first case used all aluminum heat sinks, and the data were gathered using a jet flow rate between 5-20 cubic feet/min (CFM).

Two different cases were studied, and the first case used all aluminum heat sinks, and the data were gathered using a jet flow rate between 5-20 cubic feet/min (CFM). In this case, the optimized heat sink had a thermal resistance of 0.72 ºC/W, which was lower than all of the other heat sinks except for one that had a thermal resistance of 0.6 ºC/W, although it also had a significantly higher pressure drop.A comparison of COP, was used, and it was found that the printed heat sink, , with a COP of 2.37, had a significant advantage over all of the other heat sinks that were tested. In the second case, the optimized heat sink was tested against the copper heat sinks using flow rates between 5- 15 CFM. In this case, the optimized heat sink had a lower thermal resistance than two of the copper heat sinks, but the third copper heat sink had the lowest thermal resistance, i.e., 0.249 °C/W.

In this case, the optimized heat sink had a lower thermal resistance than two of the copper heat sinks, but the third copper heat sink had the lowest thermal resistance, i.e., 0.249 °C/W. However, it was found that the thermal resistance of this copper heat sink had been decreased because its pressure drop was significantly higher than that of the printed heat sink. In this case, the optimized heat sink was again found to have the highest COP value, i.e., 0.197.Heat PipesThompson et al. used L-PBF to produce multi-layered, flat-plate oscillating heat pipes (FP-OHPs) which are typically used in high heat flux applications [8]. The FP-OHPs had a total volume of 5.08 × 3.81 × 1.58 cm3, and consisted of a closed loop, serpentine-arranged mini channel with a diameter of 1.52 mm. A fill port was attached to the side for adding/removing working fluid as necessary. Error! Reference source not found. shows a schematic and photograph of the FP-OHP.

A fill port was attached to the side for adding/removing working fluid as necessary. Error! Reference source not found. shows a schematic and photograph of the FP-OHP. Nine “vent holes” were integrated into the design to remove any powder that was trapped in the channels, which was done by repeatedly flushing the channels with acetone. Excess powder was brushed off of all exterior surfaces and the top and bottom surfaces were milled to give a smooth finish.Fig. 3. The FP-OHP with build plane (a) shown relative to fill port and (b) relative to powder bed including the laser path directions and hatch distance. Photograph of finished Ti-6Al-4V also shown [8]. Used with permission.2637The FP-OHPs were tested and found to operate successfully in both the horizontal and vertical conditions. It was found that critical power inputs were required to initiate the operation of the OHPs, i.e., between 15-20 W for the vertical operation and between 20-25 W for the horizontal orientation.

It was found that critical power inputs were required to initiate the operation of the OHPs, i.e., between 15-20 W for the vertical operation and between 20-25 W for the horizontal orientation. Effective thermal conductivities were calculated at different powers for both orientations, and found to be approximately the same at 45 °C, i.e. k = 110 W/m-K. This was roughly 500% higher than a solid block of Ti-6Al-4V, while the empty pipe had keff = 18 W/m-K.SummaryAlthough there is evidence supporting the case for using PBF to manufacture HTE, there are still significant obstacles and challenges that must be overcome and potential research avenues that have yet to be explored. For example, many of the materials and alloys typically used for HTE are not readily manufacturable using PBF, e.g. pure copper and aluminum.

For example, many of the materials and alloys typically used for HTE are not readily manufacturable using PBF, e.g. pure copper and aluminum. Furthermore, although there is much interest in designing internal features, the very nature of PBF will leave any internal feature full of powder, and removal would require either special design considerations or physical drilling into the structure and plugging the hole afterwards [8]. Furthermore, there appear to be limitations on the minimum size of any feature that can be produced due to powder size availability. However, one of the biggest opportunities for merging PBF and THE is the ability to design and manufacture highly-conformal parts with thermal management already integrated into them.

However, one of the biggest opportunities for merging PBF and THE is the ability to design and manufacture highly-conformal parts with thermal management already integrated into them. This could lead to significantly more efficient heat transfer in less space, because it would eliminate the need for external heat exchangers and completely eliminate the thermal resistance between the heat source and the heat exchanger.AcknowledgmentsThis work was sponsored in part by the National Science Foundation under Grant# ECCS 1660446.References[1] Dede, E. M., Joshi, S. N., and Zhou, F., 2015, “Topology Optimization, Additive Layer Manufacturing, and Experimental Testing of an Air-Cooled Heat Sink,” J. Mech. Des., 137(11), p. 111403.[2] Yap, C. Y., Chua, C. K., Dong, Z. L., Liu, Z. H., Zhang, D. Q., Loh, L. E., and Sing, S. L., 2015, “Review of selective laser melting: Materials and applications,” Appl. Phys. Rev., 2(4), p. 41101.[3]Frazier, W.

Mansﬁeld4, Taylor H. Pesek3, Zachary K. Snow1, Edward B. Cottiss1, Oliver D. Ebeling-Koning3, Matthew G. Price1, Mark H. Traverso1, L. David Tichnell3, Christopher B. Williams1, Alfred L. Wicks2 1Design, Research, and Education for Additive Manufacturing Systems Laboratory 2Mechatronics and Autonomous Vehicles Laboratory 3Terrestrial Robotics Engineering and Controls Laboratory Department of Mechanical Engineering 4Polymer Processing Laboratory Department of Chemical Engineering Virginia Tech REVIEWEDAbstractThe layer-wise deposition of Additive Manufacturing (AM) processes allows for signiﬁcant freedom in the design of product geometry; however, the use of 3-axis deposition tools results in layer interfaces that reduce material properties in the build direction. Adding additional degrees-of-freedom (DOF) to the AM tool could remove this limitation by enabling out of plane material deposition.

Adding additional degrees-of-freedom (DOF) to the AM tool could remove this limitation by enabling out of plane material deposition. For example, multi-DOF tool paths could align material extrusion with a part's stress contours to circumvent inter-layer delamination. As a step towards this goal, the authors designed, fabricated, and tested an AM extrusion system that leverages a 6-DOF robotic arm. In this paper, the authors detail the realization of this system including the design of a high-temperature ﬁlament extruder, kinematics and tool path generation, and user interface. The performance of the system is evaluated through layered deposition of ABS thermoplastic.Key words: Material extrusion, Additive manufacturing, 3D printing, Industrial robot, Out of plane deposition1.IntroductionMaterial extrusion is a classiﬁcation of additive manufacturing (AM) that has found In extrusion AM, a ﬂuid (usually a softened wide use by industry and hobbyists alike.

thermoplastic) is precisely extruded onto a surface. The extrusion device and/or the build surface are moved such that the extruded material forms a single printed layer. As the motion control is typically accomplished via a 3-axis Cartesian coordinate platform, parts are formed by depositing successive layers along a vertical build axis.While this method of layer-wise deposition has been shown to be capable of processing many diﬀerent materials and realizing complex geometries, it does suﬀer from some limita- tions. The layer stacking discretizes curved surfaces, which due to the large layer thicknesses common with extruding viscous melts, results in a stair-stepping eﬀect along the contour and1 1314a poor part surface ﬁnish [1, 2]. Additionally, the extrudate cools rapidly upon deposition, leaving only a short window of time in which temperatures are elevated suﬃciently for molec- ular diﬀusion across weld lines [3].

Additionally, the extrudate cools rapidly upon deposition, leaving only a short window of time in which temperatures are elevated suﬃciently for molec- ular diﬀusion across weld lines [3]. This poor weld strength results in anisotropic material properties along the build axis [2, 4]. To improve this weld strength, Partain investigated reheating the previous layer of the part just before depositing the current layer to encourage interlayer bonding [5]. Shaﬀer and coauthors added radiation sensitizers into ﬁlament and exposed the layer to radiation before depositing the next layer to promote crosslinking [6]. [7]. While Perez developed speciﬁc polymer blends that showed less anisotropic behavior these methods work for certain material types, they may not be universally eﬀective.Many groups have also looked to improve printed parts surface ﬁnish by creating adaptive slicing algorithms.

By strategically adjusting the layer thickness used at diﬀerent locations in the part (e.g., using smaller layers at places of higher curvature along the build axis), curves can be better approximated [8, 9]. Pandey et al. take a diﬀerent approach by hybridizing the material extrusion process with machining technology to smooth stair-steps created in the deposition process [10]. Curved layer slicing was developed for 3-DOF systems so that the deposition path could better follow the curvature of the part [11–13].Though means for improving the deposition strategies of material extrusion have been introduced, each of the above methods only work in very speciﬁc conditions. For instance, curved layer techniques have to be mindful of tool head and gantry collisions with the part being printed.

For instance, curved layer techniques have to be mindful of tool head and gantry collisions with the part being printed. Even with adaptive slicing, the surface is always discretized, and as the layer thickness continues to decrease, build time increases signiﬁcantly.To address prominent limitations found in extrusion AM, researchers are looking to move away from the 3-DOF gantry to more ﬂexible printing platforms such as robotic arms or CNC tables with rotary platforms. Zhang et al. document the utility of robotic arms in a number of direct metal systems [14]. These high DOF systems have already been used to great eﬀect in minimizing support material usage in Directed Energy Deposition AM processes [15–22]. Singh and Dutta have also shown that by changing the build direction throughout the part, the surface ﬁnish on a sloped or curving metal part can be improved [23]. Djuric and Urbanic have demonstrated the use of a robotic arm to deposit polymeric materials [24].

Djuric and Urbanic have demonstrated the use of a robotic arm to deposit polymeric materials [24]. Though the work restricted the deposition strategy to XY layers, the authors do note the possibilities of more novel strategies. Keating and Oxman created a material extrusion platform using a robotic arm that carried a build platform between deposition and machining areas [25]. Mounting the platform to the robot allowed the printing of overhangs without the need for support by changing the orientation of the part relative to the deposition nozzle.The authors here look to use the capabilities of a 6-DOF robotic arm to enable novel deposition strategies and process advanced composite polymers. As a step towards achiev- ing this goal, this paper describes the design, fabrication, and testing of an extrusion AM platform that employs a 6-DOF robotic arm (ABB IRB 1200-7/0.7) [26].

As a step towards achiev- ing this goal, this paper describes the design, fabrication, and testing of an extrusion AM platform that employs a 6-DOF robotic arm (ABB IRB 1200-7/0.7) [26]. Shown in Figure 1, the system features a custom built feed system and extruder, heated bed, and user inter- face. The system is built in Robot Operating Studio (ROS) [27] and accepts G-Code ﬁles from slicing software such as Slic3r [28]. The design of the system and its numerous sub-2 1315functions, including ﬁlament feed system, extrusion hot end, control software, graphical user interface, path planning, and microcontrollers, are presented in Section 2. System validation is presented in Section 3, which includes demonstrations of part printing and measurements of part strength as a function of part orientation. A discussion of the resulting system is presented in Section 4.

A discussion of the resulting system is presented in Section 4. Closure and paths for future work are presented in Section 5.Feed systemRobotic armUser interfaceHot endHeated bedFigure 1: Robotic Material Extrusion Platform2. System DesignThis section outlines the design and fabrication of each subsystem of the material extru- sion platform. The long-term goal for the platform is to develop novel deposition strategies by leveraging the capabilities of a 6-DOF robotic arm in conjunction with depositing poly- mer composite materials. No readily available slicing software has the capability to generate G-Code with Cartesian coordinates as well as tool orientation; as such new methods of slic- ing must be developed. Table 1 provides a concise list of the main requirements guiding the design of the robotic extrusion platform.The following sections detail the design and implementation of the various subsystemscomprising the platform.2.1.

Feed SystemThere are two diﬀerent types of feed materials used in material extrusion processes, pellets and ﬁlaments. As previously mentioned, the platform is targeted toward printing composite materials which have continuous ﬁber reinforcement. This reinforcement restricts the design3 1316Table 1: List of system design requirementsDesign Requirements Process 3mm filament Able to reach melt chamber temperature of 350°C Extrusion feed rate of 10 lb/hr 6-Axis (out of plane) Deposition System accepts custom .gcode Modular and expandable system architectureto a ﬁlament fed system, as the screw mechanism required to process pellets would interfere with the continuous ﬁber.Filament feed devices commonly feature a toothed gear driven by a motor, which forces feed material down the feed channel of the hot end and into a melt chamber. The completed feed system for the robotic material extrusion platform, shown in Figure 2, is based on that general concept.

The completed feed system for the robotic material extrusion platform, shown in Figure 2, is based on that general concept. The one deviation is the inclusion of tandem knobbed gears, designed to maintain grip on ﬁlament through inconsistent diameter sections. This is particularly important for the processing of the small batch novel ﬁlaments, which are still in development and often result in inconsistent diameters. Opposing the knobbed gears are spring loaded bearings, which the user can tighten using the thumb screws on the side. The hot end mounts directly to the plate below the gears.Robot mountTandem knobbed gearsSpring loaded bearingsThumb screwsHot end mountFigure 2: Isometric and front views of the assembled feed system2.2. Hot EndThe design for the hot end was based on the form found in most material extrusion applications: a heated section and a ﬁn section separated by a heat break (Figure 3).

Hot EndThe design for the hot end was based on the form found in most material extrusion applications: a heated section and a ﬁn section separated by a heat break (Figure 3). The function of the heated section is to induce phase transition in the feed material. These elevated temperatures are required to reduce the viscosity of the melt, which improves both processability and print quality. However, if heat travels up the stock material outside of4 1317the heated section, problems such as clogging the channel and buckling in the ﬁlament can occur. To prevent this, a ﬁn section was designed with forced convective cooling.Composite materials are typically more viscous than homogeneous materials and there- fore require higher processing temperatures and pressure drops. To account for this, the feed channel was designed to take 3 mm feed stock.

To account for this, the feed channel was designed to take 3 mm feed stock. The 3 mm ﬁlament will have much more resistance to buckling in the feed system than 1.75 mm ﬁlament while also allowing for a wider range of nozzle diameters to work with during testing. For initial design validation and material property tests discussed in Section 3, a nozzle diameter of 2.25 mm was chosen for a 25% reduction.The hot end was designed for printing 10 pounds of ABS an hour at 230°C with a maximum processing temperature of 350°C. In order to specify the appropriate heater size, a rough estimation of the power required to heat that output of ABS to the desired temperature was found to be 364 W. In order to achieve the maximum processing temperature, that estimate was approximately doubled. Six 100 W heaters are placed radially around the ﬁlament channel. Additionally, three thermocouples are placed up the length of the channel to monitor the temperature of the melt.

Six 100 W heaters are placed radially around the ﬁlament channel. Additionally, three thermocouples are placed up the length of the channel to monitor the temperature of the melt. The fully assembled hot end can be seen in Figure 3.Fin sectionHeat breakHeater sectionThermocouplesFigure 3: Fully assembled hot endTo validate the design as well as identify the necessary convective heat transfer coeﬃcient across the ﬁns, a thermal simulation was performed in Siemens NX 8.5. Properly specifying the coeﬃcient is necessary to prevent melting from occurring inside the ﬁn section, which can cause clogging and buckling. The simulation was performed with a 2D cross section with axial symmetry. Two of the 100 W heaters were modeled as modulated heat ﬂux sources at their respective locations in the heating block. For this simulation, a convective heat transfer coeﬃcient of 30 W m2∗K was used across the ﬁn section.

For this simulation, a convective heat transfer coeﬃcient of 30 W m2∗K was used across the ﬁn section. A set point of 230°C (the5 1318target temperature for processing ABS) was given to the bottom thermocouple with simple on-oﬀ control of the heaters. Figure 4 shows a resulting image from the simulation.Fin SectionHeat BreakHeater SectionFigure 4: Hot end thermal simulationAs seen, temperature in the melt pool reaches approximately 220°C at the top and middle thermocouple and only reaches the desired 230°C at the location of the bottom thermocouple. To remedy this, a control scheme was developed to monitor both the top and bottom thermocouples. This setup provided adequate heating for printing while also not burning the extrudate.The heat transfer coeﬃcient of 30 Wm2∗K was found to provide suﬃcient cooling across the ﬁns, appropriately dissipating the heat transferred through the heat break.

The hottest temperature found above the heat break was approximately 100°C, just below the glass transition temperature of ABS.6 13192.3. Control SoftwareThe system requires the integration of three subsystems: the main computer, robotic arm, and microcontrollers, which control the feed system motor and temperatures of the bed and hot end. The software architecture was built in ROS, which has a large open source community. Using ROS provides the ability to reuse large amounts of code from other sources, such as the ABB robot drivers and the MoveIt! motion planner.Figure 5 shows a high-level software overview. For typical use, the user would ﬁrst process the desired STL ﬁle through a slicing software to generate a G-Code ﬁle. This ﬁle then goes through a parser that pulls out movement and extrusion information. The information is converted into a tool path plan and sent to the robot.

This ﬁle then goes through a parser that pulls out movement and extrusion information. The information is converted into a tool path plan and sent to the robot. During a print, the robot is constantly tracking its own position and reporting it back to the main computer. Once the robot passes over the next point that was extracted from the G-Code, the computer sends the next extrusion command to the extruder.Figure 5: High level software ﬂow chartBuilding the system in ROS not only aﬀords the ability to recycle previously used code but also inherently provides modularity to the system through its node based architecture [27]. Each functionality of the software is written as individual communicating nodes, which allows for integration of additional functionality, such as error detection, without a full overhaul of7 1320the system. This isolation also allows for the use of diﬀerent robotic arms, as the only robot speciﬁc nodes are the ABB drivers.

This isolation also allows for the use of diﬀerent robotic arms, as the only robot speciﬁc nodes are the ABB drivers. Replacing this with a diﬀerent brand or model of arm would transfer full system functionality to that new platform.2.4. Graphical User Interface (GUI)The GUI, shown in Figure 6, is where the user calibrates the robot, locates the ﬁle for printing, slices and generates the tool path, and begins the print. It also contains controls for the feed system and heaters. Additionally, there are a number of tools available to the user for monitoring the printer including thermocouple readouts for the extruder and bed and indicators signaling successful communication.Figure 6: GUI used to control the platformThe most critical requirements for the control interface are stability and isolation. As it contains a number of safety features, the control interface cannot rely on other components of the system and it cannot crash during operation.

As it contains a number of safety features, the control interface cannot rely on other components of the system and it cannot crash during operation. The stop command available on the user interface is a safe stop for the system and will end the print. Though there are other methods of stopping the robot’s movements and extrusion built into the system (including emergency stop buttons, a dead man switch, and power breakers), these methods have the potential to cause damage to the system if used improperly or during a print.2.5. Path PlanningThere are two options for generating a tool path for the robot: the XY planar deposition process demonstrated in Section 3 and the out of plane deposition process demonstrated in Section 6.8 13211. The XY planar deposition path planning operates exactly the same as most commercial printers. The user imports the desired STL ﬁle through the GUI and slices it, which produces a G-Code ﬁle.

The XY planar deposition path planning operates exactly the same as most commercial printers. The user imports the desired STL ﬁle through the GUI and slices it, which produces a G-Code ﬁle. This can be done using a number of diﬀerent slicers, but it [28]. The user then selects to plan the tool path is currently operating using Slic3r which performs all of the necessary inverse kinematics before printing can begin.2. Out of plane deposition requires rotation information in addition to Cartesian coordi- nates. As such, a standard slicer is insuﬃcient. To generate this tool path, the part is sliced ﬂat on the bed as though it would be printed traditionally. The resulting G-Code and the desired part orientation is then passed to a custom MATLAB script that takes the orientation information, rotates the previously generated G-Code and creates a corresponding quaternion. The quaternion is then added to each move com- mand in the G-Code.

The quaternion is then added to each move com- mand in the G-Code. This new G-Code ﬁle can then be imported to the GUI where the user can initiate the same path planning algorithms used in the XY path planning approach.2.6. MicrocontrollerThe designed system uses two microcontrollers to communicate between the end eﬀector, the heated bed, and the main computer. The system required communication over TCP/IP, SPI, CAN, and I/O switching capabilities, as such, the TI TIVA TM4C1294XL microcon- Interfacing with the microcontrollers were a number of peripherals: troller was selected. K-type thermocouples, EPOS 2 motor controller, and heater relay control. The thermocou- ples are used to measure the hot end melt chamber temperatures in three locations and the heated bed in two locations. The EPOS 2 motor controller drives the Maxon motor used in the feed system. The relays control the power to the heated bed and hot end.

The EPOS 2 motor controller drives the Maxon motor used in the feed system. The relays control the power to the heated bed and hot end. As there were no existing boards to interface with the microcontroller and desired peripherals, a custom shield, shown in Figure 7, was designed as a compact solution.Figure 7: Custom shield for the TIVA microcontroller as designed (left) and produced (right)Basic temperature control is achieved by switching the relay on and oﬀ around maximum and minimum temperature set points. As an additional safety feature on the extruder side,9 1322another relay is used as an enable relay in conjunction with the main heat control relay. This is in response to relays being mechanical devices with a limited number of switching cycles. If the control relay broke closed during a print, the uncontrolled temperature rise could potentially damage the hot end.

If the control relay broke closed during a print, the uncontrolled temperature rise could potentially damage the hot end. The enabling relay switches on when ﬁrst turning on the heaters and will switch oﬀ if any anomalies are detected in the switching of the control relay.3. System Validation3.1. Base FunctionalityInitial tests were performed in stages. After communication between the computer, microcontrollers, and robot was established, basic extrusion tests without robot movement were performed. This was to ensure proper heating of the ﬁlament and suﬃcient cooling across the ﬁns. Figure 8 shows temperature data from the top hot end thermocouple during the simple extrusion test with a set point of 240°C.Figure 8: Steady state temperature data of the top hot end thermocouple while printingThe temperature creates a periodic pattern around 250°C where the heaters are switched on when drifting slightly below the set point, overshooting, and then ﬁnally cooling down.

This is due to the distance between the heaters and thermocouples, which causes the heaters to heat the area around them faster than the temperature rise registers at the thermocou- ples. A more advanced control scheme that accounts for the delay has been developed but not tested. Next, basic print parameters, shown in Table 2, were found empirically. This conﬁguration was used to create the sample prints shown in Figure 9.10 1323Table 2: A list of the most relevant print parameters used for making the tensile specimensParameter NameLayer Height Perimeter Speed Infill Speed Non-print Speed Bed Temperature Extruder Temperature Skirt Loops1.4 mm 10 mm/s 20 mm/s 40 mm/s 100°C 235°C 2 (1 for Z)3 in5.5 inFigure 9: Sample prints from the robotic extrusion platform11 13243.2. Mechanical Property Testing3.2.1. Testing MethodsSystem performance was veriﬁed through tensile strength tests on directly printed spec- imens in accordance with ASTM D638 [29].

Mechanical Property Testing3.2.1. Testing MethodsSystem performance was veriﬁed through tensile strength tests on directly printed spec- imens in accordance with ASTM D638 [29]. Three tensile specimens, each with an overall length of 165 mm, were printed with the robotic material extrusion system in each Cartesian direction using ABS plastic. All specimens were printed with inﬁll roads strictly in the X direction. The ﬁrst layer of each specimen type is shown in Figure 10. An extruded “skirt” was used to clear out the nozzle to ensure the quality of extrudate during printing. The spec- imens were created using the previously established print parameters.

An extruded “skirt” was used to clear out the nozzle to ensure the quality of extrudate during printing. The spec- imens were created using the previously established print parameters. System calibration and material spool were also kept constant throughout the test.SkirtContourYInfillXFigure 10: The ﬁrst layer of each specimen type: X (left) Y (center) Z (right)Due to the small cross sectional area of the Z specimens, the recently ﬁnished layer and those below it were still soft when beginning the next layer. As these layers continued to stack up, the specimen lost all shape and stability, causing the build to quickly collapse. To prevent this, a delay of 20 seconds was placed after each layer to allow the specimen to suﬃciently cool for the subsequent layer. This delay allowed the nozzle to ooze ﬁlament, compromising the beginning of the next layer.

This delay allowed the nozzle to ooze ﬁlament, compromising the beginning of the next layer. In response, the same skirt used in the ﬁrst layer was placed at the beginning of each layer to allow the nozzle to clear itself. This had the added beneﬁt of creating additional support structure, shown in Figure 11, for the specimen as it grew. The thickness of the specimen was also reduced in order to eliminate the need for inﬁll, as the inﬁll roads were also found to compromise the quality of the specimen.The Z specimen did require post processing in the form of removal from the stacked skirts. The threads connecting it to the skirts were cut and excess material threads were removed. The X and Y specimens were able to be tested as printed with no additional processing. Figure 12 shows a printed specimen of each type.12 1325Figure 11: Z specimen as printed.

The X and Y specimens were able to be tested as printed with no additional processing. Figure 12 shows a printed specimen of each type.12 1325Figure 11: Z specimen as printed. The stacked skirts allow the nozzle to clear while also supporting the specimenFigure 12: Specimens as printed, each with an overall length of 165 mm. X (top) Y (middle) Z (bottom)13 13263.2.2. Testing ResultsThe tensile tests were performed on an Instron 5894 and followed ASTM D638 [29]. The Z specimens broke too quickly for proper measurement with the standard 5 mm/min extension rate, so a rate of 2 mm/min was used for those tests. An extensometer was attached to the gauge section of each specimen during testing. Stress-strain curves for all specimens are shown in Figure 13.Figure 13: Stress-strain curves for all tensile specimensThe modulus was calculated using a linear ﬁt through the data set. The amount of data used in the linear ﬁt was reduced until the R squared value exceeded 0.995.

The amount of data used in the linear ﬁt was reduced until the R squared value exceeded 0.995. Yield strengths were calculated using the Oﬀset Yield Strength method described in ASTM D638 with an oﬀset of 0.2%, the results of which are shown in Figure 14 [29]. The “aligned” data set presents data from samples similar to the X specimen set with roads aligned down the length of the tensile specimen [30].14 1327ModulusYield Strength300030250025200020) a p M(s u l u d o M1500AlignedXYZ) a p Mh t g n e r t Sd l e i Y15100010500500Figure 14: Moduli and yield strength results4. DiscussionThe results of the tensile tests were as expected. The in-plane strength is much greater than that of the out of plane strength. The one discrepancy from expectation is the better performance of the Y specimens over the X. This is a result of the following:1. The layers in the Y specimens were deposited more quickly than the X specimens.

This is a result of the following:1. The layers in the Y specimens were deposited more quickly than the X specimens. The tool path for the X specimens was such that a number of travel (non-printing) moves were required to complete a layer. These travel moves add additional time per layer, allowing the part to cool, which limits interlayer diﬀusion during deposition. This is in contrast with the tool path for the Y specimen, which only had two travel moves per layer: a move from the end of the contour to the start of the inﬁll and a move from the end of the inﬁll to the start of the next contour.2. The inﬁll pattern of the Y specimens reinforced the contour. During deposition of the contour for each layer, a small gap was occasionally created where the start and end points of the extrusion did not overlap, creating a weak point. This event occurred in all specimens, but the Y specimens were able to compensate for the gaps with inﬁll.3.

This event occurred in all specimens, but the Y specimens were able to compensate for the gaps with inﬁll.3. The inﬁll pattern consisted of many short roads. This allowed for enhanced bonding between the roads in the Y specimens over the X specimens, as those roads were much longer and therefore had additional time to cool before an adjacent road was placed. A part will not necessarily have exclusively short roads, so the performance of the Y specimens may be inﬂated.These concerns all pertain to the way in which the specimens were produced. By keeping the fabrication method (i.e. the tool path) the same between the X and Y specimen sets, these issues would not appear. As such, future testing should examine these same properties from samples machined from a printed sheet.15 1328AlignedXYZAn additional point of note is the inconsistency in the Z specimens.

Figure 14 shows that the yield strength standard deviation is 2.5 times greater in the Z direction than in the X direction. This variation results from two causes:1. The printer resolution. The XY cross sectional area of the Z specimens (3.49 mm by 19 mm in the grip section) is close to the nozzle diameter (2.25 mm) of the printer, thus requiring only two passes per layer. The tool path nominally takes the form of a rectangle, but as the robot makes the sharp corners at each end the cross section is rounded into an oval.2. The fabrication method. The road cross-sectional area is much larger on this platform than those from a traditional material extrusion system. As such, they retain heat longer. Though this behavior is beneﬁcial for inter-road bonding, it also compromises print quality for small prints such as this. The artiﬁcial delays between layers were necessary to have a survivable sample, but greatly reduce the potential for layer bond- ing.

The artiﬁcial delays between layers were necessary to have a survivable sample, but greatly reduce the potential for layer bond- ing. That said, the delays were also not long enough for the previous layer to fully cool, causing it to deform during the deposition of the next layer. This further exacerbates the resolution issue by warping the cross section.Both of these issues are prevented with a larger layer. The print time per layer increases and the resolution issues caused by the large nozzle are diminished with larger travel times. As such, machining the Z specimens from a block of material rather than directly printing them may give a better estimate of the material properties one could expect from a print.5. Conclusion and Future WorkThis paper detailed the design, fabrication, and testing of a robotic material extrusion platform. The custom extruder and feed system are shown to be capable of processing ABS plastic.

The custom extruder and feed system are shown to be capable of processing ABS plastic. The system is capable of processing standard G-Code ﬁles output from slicing software and realizing complex structures such as those seen in Figure 9. Material tests have also shown that the printed ABS is, expectedly, anisotropic with an in-plane yield strength of 20 MPa, 50% of the nominal strength.The presented platform is designed to extend the capabilities of typical 3-DOF AM extrusion by leveraging a 6-DOF robotic arm. The additional DOF provide the ability to rotate the tool head, allowing for out of plane deposition. An example of this can be seen in Figure 15.

The additional DOF provide the ability to rotate the tool head, allowing for out of plane deposition. An example of this can be seen in Figure 15. Sacriﬁcial scaﬀolding structures were deposited in the XY plane, but the model material was deposited at an angle, normal to the surface of the desired part.16 1329Figure 15: Tensile specimen printed using out of plane deposition techniquesThis idea can, as previously established, eliminate the stair-stepping eﬀect in a number of parts, but also improve material properties. As seen, the tool path of the desired object resembles that of the X specimens discussed in this paper. This is opposed to if the object was printed with a 3-DOF gantry, where the tool path would much more closely resemble that of the Z specimens. As discussed, the X specimens dramatically outperformed the Z specimens. The authors feel there is potential for similar behavior with out of plane deposition.References[1] M. K. Agarwala, V. R. Jamalabad, N. A.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDESIGN AND ROBOTIC FABRICATION OF 3D PRINTED MOULDS FOR COMPOSITESRajkumar Velu*, Nahaad Vaheed, and Felix Raspall*Digital Manufacturing and Design Centre, Singapore University of Technology and Design, Singapore 487372 *Corresponding Authors: felix_raspall@sutd.edu.sg, rajkumavelu@sutd.edu.sgAbstract3D printing technologies have a direct impact on manufacturing the composite structures and in particularly fabrication of molds. Molds produced through additive manufacturing methods would greatly improve product features. The material selection and process conditions involved for producing mold tooling, mainly towards Automated fiber placement (AFP) work cells.

The material selection and process conditions involved for producing mold tooling, mainly towards Automated fiber placement (AFP) work cells. In this study, the main objective is to improve the design and fabrication of composite parts through complex molds as well as to assess and improve the production workflow through the development of an effective design environment for the existing fiber placement operation. A robotic arm will be used to hold the print surface and to follow a pre-programmed print path with a stationary extruder to fabricate the mold tooling. This paper will present a review on the selection process for mold materials and the initial experimental work carried out to investigate required properties of 3D printed molds.IntroductionMolds are fabricated using composites also known as tools, which is plausible to virtually fabricate for any materials. The molds are in the form of blocks or series of blocks or tailored into required shape for fabrication process.

The molds are in the form of blocks or series of blocks or tailored into required shape for fabrication process. The fabrication involved in different methods such as casting or pliable raw material using rigid frame or mold called lamination method as shown in figure 1 respectively. In the casting method the materials are melted in to liquid form and poured into the mold for curing to form a required shape as solid state [1]. In the lamination method the prepreg sheets and fiber reinforced composite materials are laminated on the mold material to form a required shape [2&3]. This is a form that mimics the final shape of the part, whereas the part is constructed over its outer surface which also called as positive mold. The positive type of mold is quiet rapid fabrication process relatively. The molding techniques are classified in to different procedures based on specific purpose and materials.

The positive type of mold is quiet rapid fabrication process relatively. The molding techniques are classified in to different procedures based on specific purpose and materials. The noteworthy techniques are vacuum forming, blow forming, compression molding, injection molding and laminating [4-8]. In this current study focused the mold fabrication for prepreg lamination techniques.1036Metal moldComposite laminatesMoldCasting(a) Casting method(b) Laminate methodFigure 1 Fabrication of parts or laminates using distinct types of moldsFurther the composite tools can classify in to two types such as hard tooling and soft tooling, which is based on type of material and process used. Ceramic and metal tools are usually used as hard tools and these materials can withstand thousands of production cycles for same set of design [9].

Ceramic and metal tools are usually used as hard tools and these materials can withstand thousands of production cycles for same set of design [9]. Also, there are excessive cost and performance tool material such as steel alloys called by Invar and it requires regular services with a specialist. When comes to soft tooling, they are quite easy to fabricate, because of fabricating with the materials like those composite manufacturers will use for the part. Subsequently, a significant key factor with tooling for composites is the phenomenon of coefficient of thermal expansion (CTE) [10]. In general, the most common materials choices for composite tools are steel and aluminum due to low cost and high-performance, but they mismatch with composites CTE. Thus, affects the quality of the dimension stability of final composite structures.

Thus, affects the quality of the dimension stability of final composite structures. If the materials like metal alloys such as Invar are close to composite CTE, which increase the dimensional accuracy during cure because the shrinkage and thermal expansion of the tool and part will be similar. Subsequently the other most common molds are made of rubber or other flexible materials, whereas most casts are made of rigid-setting materials, although this is not always the case [11]. Some materials can be used for both molds and casts and are available in a variety of flexibility and firmness properties. Among the most commonly used mold materials are liquid latex, silicone rummer, urethane rubber, and alginates.

Among the most commonly used mold materials are liquid latex, silicone rummer, urethane rubber, and alginates. Common cast making materials include Plaster-of-Paris and Gypsum cement, concrete, plastics (resins and epoxies), waxes, metals [12 &13]However apart from selection of materials, the fabrication of the mold will be one of the most important key factors among the physical properties of the mold, cost of fabrication and duration to construct the mold [14]. Further, the number of parts to be made with a mold, mold service life, the final product specifications, tolerance and surface finish also the noteworthy factors for fabricating the molds as shown in figure 2. There are different techniques or machining process to fabricate the mold which all based on the specific products. As mentioned earlier, in this study focuses on mold fabrication for prepreg composite structures.

As mentioned earlier, in this study focuses on mold fabrication for prepreg composite structures. The prepreg layup on substrate or mold are classified into two types they are Automated tape placement (ATP) [15] and Automated fiber placement (AFP) [16&17]. Both the processes are involved similar functions; however, each1037method is used for different specific structure of fabrications based on mechanical strength or stiffness needed. These processes apply resin-impregnated continuous fiber. Among these techniques AFP process has unique features and automatically places multiple individual preimpregnated tows onto a mandrel at high speed., using a numerically controlled placement head to dispense, clamp, cut and restart each tow during placement. Minimum cut length (the shortest tow length a machine can lay down) is the essential ply-shape determinant [18].

Minimum cut length (the shortest tow length a machine can lay down) is the essential ply-shape determinant [18]. The fibre placement head can be attached to an existing gantry system, retrofitted to a filament winding machine or delivered as a turnkey custom system.Maintenace and recyclingDimensional accuracyMaterial selectionMold durabilityFactors involved for fabrication of mold/toolsProduct specificationTime consumptionFabrication costFigure 2 Factors involved for fabrication of mold/toolsAFP process achieves high production rate, better quality, efficient and low cost of manufacturing of large scale composite structures. When integrating AFP process with the robots, leads to highly automated process, further reduction in material wastage, good processing quality and repeatability. Also, the product outcome from AFP process has significant dependency on mold material property and surface.

Also, the product outcome from AFP process has significant dependency on mold material property and surface. The conventional mold making process for AFP techniques are CNC machining and hand crafting. Both methods are used in all industries, for small fabrication parts handmade tools are more prevalent whereas no need of complicated CNC processing [19]. In general, the handmade mold is fabricated similarly like composite part itself, by layering resin – impregnated fabrics on top of a master model, curing the mold and removing the master to produce the cavity. This process involves labor intensive process requiring skilled workers, and to meet the required mold design specification, will undergo for various iterations which leads to substantial time consuming. Correspondingly, the mold tools fabricated on CNC machining has certain limitations like significant labour and machining, leading to excessive costs, significant material1038waste, and long lead times.

The materials mostly used for machining process are metallic materials (aluminium, steel, or Invar alloys) [20], although non-metallic materials are also frequently used, including specialized fibre-reinforced polymer (FRP) materials, high-temperature tooling board, and others [21]. The use of metal or other hard tooling drives the need for highly complex, collapsible designs. Inflatable bladders require investment in additional tooling, adding cost and time, and are limited in the geometries.To overcome the aforementioned limitations using conventional fabrication of mold such as CNC or handmade crafting, the manufacturing industries propelled towards Additive manufacturing or 3D printing techniques for fabricating molds or substrate which is also called by Rapid Tooling (RT) [22 &23].

This method incorporates methods of fabricating the objects by building through additional of materials, significantly opposing subtractive methods such as CNC machining, which remove materials until achieving a final shape. Fabrication of composite structures is the most original forms of additive manufacturing techniques. There are distinct types of lay-up process involved for fabricating composite structures are distinctly additive in nature, building layer by layer to form final parts. These advantages of additive manufacturing have been revolutionized with the advent of the 3D printing industry [24]. Thus, the reason the manufacturing industry propelled towards 3D printing techniques due to more pressure on global competition.

The significant advantages of the 3D printing techniques for fabrication of molds towards conventional techniques are low cost, complicated design possibilities, reduction in material wastage where in CNC operation involves milling, or sanding requirements, automated manufacturing, ability to recycle waste material, minimal inventory risk, improvement in capital management, ability to easily share designs and outsource manufacturing, speed and ease of designing and modifying products as shown in figure 3.Design freedomImprovement in capital managementSpeed and ease of Designing and modificationMinimal inventory riskFabrication of 3D mold/toolsLow Material wastageRecycle of waste materialAutomated manufacturingFigure 3 Advantages of 3D printed mold fabrication1039There are different varieties of Additive manufacturing technologies and they all have some potential relevance for manufacturing the composite structures [25], however, in this study focused on extrusion-based technologies, as it is current relevance and a clear path to significant future impact as well [26].

The extrusion-based technology also known as Fused Deposition modelling patented by Stratasys [27]. In this method the parts build layer-by-layer by heating and extruding thermoplastic materials in an optimized controlled parameter. In general, the FDM system that are applied in industry or research sectors are equipped with three linear motion axes in order to move the process-specific tool head relatively to build-up platform in a translational motion. Based on the CAD model, a slicing program segments the geometry and plans the process paths for each slice. While slicing and modelling strategy plans planar, parallel layers that are aligned horizontally and that leads to supporting structures for geometry overhangs depending on the determined build-up direction. To break through this modelling paradigm, a multi-axis kinematic to guide the process which is called as robot-based extrusion method [28].

To break through this modelling paradigm, a multi-axis kinematic to guide the process which is called as robot-based extrusion method [28]. Integrating the additive manufacturing design freedom method and the multi-axis control or industrial robotics are further step towards development in the 3D printing revolution. The complex design fabrication can be achieved, fiber alignment can be manipulated and controlled degree also reliable when adding multi-axis motion platforms. The X-Y gantry system has limitations on complex structure lay-up, transfer molding, filament winding and automation.

The X-Y gantry system has limitations on complex structure lay-up, transfer molding, filament winding and automation. The robot-based 3D printed mold has unique features like potential on fabricating light weight structures, freedom on degrees of geometric complexity, part consolidation and design optimization.Rieger et al [29] studied and investigated the application concept of additive manufactured dies to support the robot-based incremental sheet metal forming process (Robot-forming) to produce sheet metal components in small batch sizes. The result proved that AM modelled dies can indeed stabilize the sheet to reduce the deformation during the forming process compared to conventional incremental sheet forming. Viridis3D commercialized the first robotic 3D printer, the RAM 123 and achieved patent on four-axis ABB IRB460 robot to create sand molds and cores for casting metal parts.

Viridis3D commercialized the first robotic 3D printer, the RAM 123 and achieved patent on four-axis ABB IRB460 robot to create sand molds and cores for casting metal parts. The robot is equipped with as powdered material feeder that distributes the sand and a print head that dispenses the liquid binder into the sand. Spreading sand and dispensing binder intermittently, the robotic 3D printer builds the mold layer by layer. The molds are built on a stationary table. The tabletop is a pallet that can be used to move parts on and off the machine with a forklift. the main advantage is they can quote parts with a short lead time. Molds and cores with complicated shapes can be made without a lot of tooling. Going digital is also easier. Viridis3D customer Trident Alloys sees the future in this robotic additive manufacturing system. The foregoing literature review and discussion clearly indicate that 3D printed mold break throughs with significant advantage.

The foregoing literature review and discussion clearly indicate that 3D printed mold break throughs with significant advantage. Significantly, it becomes important to establish the design and development of robotic 3D printed molds, whereas the research gap identified is the development of 3D printed mold for AFP process is not yet been studied or investigated systematically so far. This paper presents a robotic additive manufacturing process for mold fabrication and introducing the integration of AM and AFP process. Which pushes one step further in the direction of optimizing complex part design and manufacturing with building parameters in consideration of robotic AM and AFP process.1040Methodology or statement of workThe research is structured into four components such as design of composite parts, Toolpath planning and workcell setup as shown in figure 4. The design of composite components is iterative in nature and requires numerous steps and tools.

The design of composite components is iterative in nature and requires numerous steps and tools. Succinctly, it covers geometric design, fiber distribution, structural analysis and structure and form optimization. These different steps require specialized software tools and designers are often converting information from tool to tool. Further in the tool path planning, the automation of the manufacturing process involves the development of processes that translate the designed part and fiber layout into manufacturing sequencing and specific machine instructions. Due to the complexity of the fabrication process, including two robotic arms with specialized end-effectors, the toolpath planning requires careful coordination, motion and collision simulation. The research developed with the required systems to optimize and streamline the toolpath planning process.

The research developed with the required systems to optimize and streamline the toolpath planning process. In this way, the fiber placement process becomes a part, or option, in a more ambitious production line that includes additive manufacturing and subtractive manufacturing processes. The end-goal is to have an operative cell capable of producing functional, high performance composite parts. Numerous industries take advantage of the high mechanical performance, low-weight of composites. Aeronautics leads the research with requirements for expensive, low volume, high-performance parts. However, as the accessibility of composite parts grows, the market can expand. This fourth theme of the research will focus on how to expand the use of composites in existing and alternative markets as lower design and manufacturing barriers can provide composites with an edge.

This fourth theme of the research will focus on how to expand the use of composites in existing and alternative markets as lower design and manufacturing barriers can provide composites with an edge. Preliminarily, we are looking at high performance, medium size parts (between 20 and 100cm long, three dimensional parts) such as UAV, electronic chassis, furniture, and architectural connectors, among others.Design of composite partsTool path planningApplication and prototypeWork cell setupFigure 4 Process flow of manufacturing process1041Novel approach and AM Robust work cell designThe robot-based 3D composites are adopted to increase the potential and the flexibility of the manufacturing process. AM robust work cell system is proposed to produce geometric specific molds that support the AFP process for composite fabrication. The modelled mold represents the negative form of the target geometry.

The modelled mold represents the negative form of the target geometry. In the AFP process, the AFP layup tool head forms the prepreg tape towards the mold shape in order to ensure the accurate position of the prepreg sheet. Figure 5 shows the bird view of our new concept work cell layout, to produce process- and geometry-specific mold sequentially to the prepreg layup process steps.Automated fibre placementprocess robot work cellAdditive Manufacturingrobot work cellFigure 5 The bird view of our new concept robotic work cell layout using AM and AFP process techniquesThe additive manufacturing robot fabricates the mold structure and the AFP robot places the prepreg tapes on the surface of the substrate. Eventually, the work flow would be continuous process like the AM robot arm will move on to AFP robot for tape placement after the substrate have been build using extruder on the work platform. The AM work cell is designed as shown in1042figure 6.

The AM work cell is designed as shown in1042figure 6. The build platform is designed and fitted at the end of the robot arm and the extruder placed above the platform in the aluminum frame. In general, the extruder moves in X and Y direction and the build platform moves along Z direction in the FDM system, here the concept focused to fix the extruder and the robot arm fixed with build platform will move in all direction with respect to the part design. Because, once the mold is produced the arm will move onto AFP side for prepreg placement process. The build platform is designed and analyzed to withstand the working load. The build platform is made up of aluminum frame and wooden platform is fixed on top.

The build platform is designed and analyzed to withstand the working load. The build platform is made up of aluminum frame and wooden platform is fixed on top. Also, the heating element is also incorporated with the building platform to maintain the glass transition temperature or just below melting temperature of working polymers to achieve the better dimensional stability.Figure 6 Additive Manufacturing robot work cell designThe concept focuses as the FDM process that melts a thermoplastic filament and extrudes a pellet to model the mold structure. The major advantage of the robot-based approach is the possibility to produce geometry-specific mold adapted to the target shapes at different forming stages. The process chain to produce is digitally defined and no geometry-specific tool is required.

The process chain to produce is digitally defined and no geometry-specific tool is required. Also, the robot manufacturing method makes the possibility to realize the modelling process in free determinable directions.1043Initial experimentsThe evaluation of the processing parameters is divided into two segments. They are AM and AFP process. Though initially the research begins with AM process alone to test the additive manufacturing mold with varied sizes. Therefore, the initial design parts are produced by means of AM robot system as shown in figure 7, since the described robot-based FDM process, and the tool path are generated based on the shape of the mold part structure. The initial experimental validation for the extrusion process is performed on AM robot. The experimental work influences the robot moving speed, extruder diameter and melting temperature. After certain trials, robot constantly moves at a speed of 7mm/min.

The experimental work influences the robot moving speed, extruder diameter and melting temperature. After certain trials, robot constantly moves at a speed of 7mm/min. all experiments are performed with 3mm of nozzle diameter with PLA and ABS material. The working temperature adjusted based on the material used during printing. In general ABS material has exemplary thermoplastic characteristics, which is used to transfer the results to the group of thermoplastic extrusion materials.Figure 7 Test specimen produced using AM robot systemConclusion and Future workAs described above, this paper initially overviews the need of mold fabrication technology and converging into the additive manufacturing technology roles in fabricating the mold for Automated fiber placement process. Which briefly describes the conventional fabrication method and its limitations. Then introduces robotic additive manufacturing concept and its requirements in the advanced manufacturing method.

Which briefly describes the conventional fabrication method and its limitations. Then introduces robotic additive manufacturing concept and its requirements in the advanced manufacturing method. Based on this new novel approach AM robot system, the initial trials evidence the general feasibility and the potential of mold modelling using robot-based extrusion process. The initial experiments have shown that AM modelled mold can indeed stabilize the prepreg sheets while fabrication. From the above description and discussion, the conclusions and future works are: Robotic additive manufacturing has a unique advantage in complex geometry printing and large-scale manufacturing. It can be simulated in an industrial robot simulation and1044offline programming software platform. The future work is to model complex geometry modelling mold to fabricate the composite structures.

The future work is to model complex geometry modelling mold to fabricate the composite structures. In addition to develop the system that uses the process simulation as a tool in AM part design and optimization with respect to the mechanical, thermal and electrical property requirements. To develop the large structures mainly focusing on aerospace and medical applications.References1. McDavid, R. M., & Thomas, B. G. (1996). Flow and thermal behavior of the top surface flux/powder layers in continuous casting molds. Metallurgical and materials transactions B, 27(4), 672-685.2. Young, W. A., Rupel, K., Han, K., Lee, L. J., & Liou, M. J. (1991). Analysis of resin injection molding in molds with preplaced fiber mats. II: Numerical simulation and experiments of mold filling. Polymer composites, 12(1), 30-38.3. Throne, J. L. (2002). Thermoforming. Encyclopedia of Polymer Science and Technology. 4. Taylor, C. A., DeLorenzi, H. G., & Kazmer, D. O. (1992).

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDESIGN AND SIMULATION OF 3D PRINTED AIR-COOLED HEAT EXCHANGERSR. A. Felber*, N. Rudolph*, and G. F. Nellis*Department of Mechanical Engineering, University of Wisconsin, Madison, WI 53706AbstractThe use of material extrusion with conductive fillers is explored for air-cooled heat exchangers. A general overview of the manufacturing tasks, design criteria, printability constraints, and modeling techniques is given, along with experimental data from prototype testing. The first sub-scale prototype design is an air-water crossflow heat exchanger designed to transfer around 100 Watts. It was printed with unfilled conventional ABS and the air channels designed with an array of round pin fins to enhance heat transfer.

It was printed with unfilled conventional ABS and the air channels designed with an array of round pin fins to enhance heat transfer. The prototype was also CT- scanned for inspection of the printed pin fin shapes.IntroductionDry-cooling reduces or eliminates the consumption of water, which has both ecological and economic benefits. Power plants would have greater siting options if they did not need a large local water supply, thus relieving the limitation that arid regions cannot host power generation sites. Current drawbacks associated with implementing dry-cooled systems include poor efficiency due to warmer heat rejection temperatures and high capital costs associated with conventional dry-cooling systems that consist of large arrays of metal finned-tubes and plate-fins.Advances in additive manufacturing, commonly referred to as 3D printing, have redefined the theoretical manufacturing limitations across a variety of applications.

This project examines the use of 3D printing in the design of heat exchanger systems for dry cooling. Additive manufacturing techniques have a range of printing capabilities, material properties, and material costs. The technique that is most generally associated with 3D printing is material extrusion, otherwise known under the trademark Fused Deposition Modeling (FDM®). For material extrusion, raw solid polymer filament is melted and extruded through a CNC nozzle. This technique has the advantages of efficiently producing functional and durable parts with complex geometries using relatively inexpensive material. The low cost of the material and the manufacturing process allow it to be considered for large-scale systems such as those required for a dry-cooled power plant.One characteristic of the material extrusion process is its range of material compatibility.

While polymer filaments have the advantages of being relatively low cost, lightweight, durable, and corrosion-resistant, their thermal conductivity is 10 to 100 times lower than that of metals (Joen et al., 2009). This is an obvious disadvantage for their use in a heat exchanger application. However, some studies have found that replicating a conventional tubular heat exchanger with polymer decreased its cost by 2.5 times, despite its larger size, because of its decreased density and material cost (Zaheed and Jachuck, 2004). Further improvements could be made with higher conductivity and robust printability. Current work, including work associated with this project, is2250being done to increase the thermal conductivity of these polymers by embedding conductive fillers such as graphite, carbon black, carbon fiber, or metal particles of various shapes and sizes into the polymer filament.

The use of filled polymers could increase the thermal conductivity by a factor of 10 or more (LATI, 2013).Overview of Manufacturing TasksThe filled polymer material used for the manufacturing in this project is not commercially available. Its thermal and mechanical properties are dependent on its filler type, volumetric content, and fabrication technique, and can be strategically designed to create unique and advantageous characteristics for the application of interest. Traditional material extrusion printing equipment is also not necessarily suitable for all filament types by default.

Traditional material extrusion printing equipment is also not necessarily suitable for all filament types by default. The manufacturing efforts associated with the fabrication of the heat exchangers include compounding of the material, filament extrusion and characterization, and modification of the conventional material extrusion process for thermally conductive filament.Compounding material and filament fabrication require unique process settings and conditions to achieve desired homogeneity and compound properties. It is especially important to ensure the material is extruded with a uniform diameter and is flexible enough to be wound into spools. Doing so with unfilled nylon material has its own difficulties, and doing so with filled nylon, embedded with conductive fillers, makes the task uniquely challenging.Assessment of the filament also plays a role in its fabrication.

For both manufacturing and design purposes, it is important that measuring and controlling material properties is also investigated in the fabrication process. Depending on the filler material, specifically the type and shape, some material properties may also be different in axial and radial directions.Once the raw material is compounded and ready for printing, some changes are required to the material extrusion process. The incorporation of fillers increases the stiffness and viscosity and requires modification of equipment for combatting these changes. Specifically, stagnation and clogging can be prevented by redesigning the nozzle with a more gradual taper.DesignThe design of the heat exchanger puts the air and water channels in cross-flow. This is typical of most gas-to-liquid heat exchangers where more of the surface area is dispersed by a large frontal area rather than long gas-side channels in order to conserve pressure drop.

This is typical of most gas-to-liquid heat exchangers where more of the surface area is dispersed by a large frontal area rather than long gas-side channels in order to conserve pressure drop. The specification of the heat exchanger geometry is broken into two categories: the microstructure, which refers to dimensions of an individual channel that can be repeated indefinitely, and the macrostructure, which refers to the overall size of heat exchanger formed by repeating and lengthening channels as necessary. The macrostructure depends on the microstructure geometry, the number of rows, and the number of air channels in each row. The parameters that define the heat exchanger macrostructure (Figure 1) include the following: height of the heat exchanger (defined by the number of rows), width of the heat exchanger (defined by the number of air channels in each row), and the depth of the heat exchanger.2251STDthwallSLhtwthfinFigure 1: Geometric configuration.

Left: air-side microstructure. Right: macrostructure.Rectangular ducts were chosen for both the air and water channels. It was recognized that a larger number of shorter water channels would be more advantageous than a smaller number of taller water channels to maximize efficiency and surface area. For this reason, each water channel was sized to have a fixed 1-mm height and its width would be adjusted with the macrostructure.The air ducts are filled with a staggered array of circular cross-sectional pin fins, as shown in Figure 1 on the left side where the top of the duct is transparent. The geometric parameters required to define this microstructure include: air-side channel height (ht) and width (w), the thickness of the walls separating the air channels (thfin), the thickness of the walls separating the water and air channels (thwall), the diameter of the pin fins (D), and their spacing in both the transverse (ST) and longitudinal (SL) directions.

Straight circular pin fins were chosen for the initial design because their performance can be predicted using existing flow correlations associated with extensive research and literature for flow over tube banks. These existing correlations enabled immediate analysis of this geometry which allowed parallel progress to be made on the manufacturing side of the project.It is understood, however, that circular pin fins may not be the optimal design, and alternative geometries are possible using material extrusion. Current work is being done to investigate different shapes of these pin fins that go beyond straight circular cylinders. For example, by streamlining the cross-section of the pin fins, the drag force can be significantly reduced, resulting in a lower pressure drop across the array. One study has been carried out specifically for flow over various cross-sectional shapes including elliptic, drop-like, and airfoil cross-sections for long fins (Sahiti et al., 2006).

One study has been carried out specifically for flow over various cross-sectional shapes including elliptic, drop-like, and airfoil cross-sections for long fins (Sahiti et al., 2006). Another improvement being investigated is the effect of varying the cross-section of the pin fin along its length such that they have larger cross- sectional area at the base, where the rate of conduction is highest, than they do at the center of the duct, where conduction is zero. Additionally, unconventional techniques for design are being used; see the section titled “Optimization”.2252Printability ConsiderationsThe design of the prototype to be printed and tested has a few additional considerations to be made regarding its printability, which mostly stems from printing parameters, one of which is the print orientation. The print orientation greatly influences how the part will be printed.

The print orientation greatly influences how the part will be printed. It refers to the direction in which the part is built, which dictates how the part is sliced into cross-sections. In most cases, a certain print orientation is chosen to minimize print time and support structure. Since both of these factors affect the cost of the print with increased labor, machine use, and material required, the orientation is not a trivial parameter.A general starting point was to set the side of the part with the largest surface area parallel to the build platform (Figure 2) because it is more efficient to print a smaller number of large cross- sections than a larger number of small cross-sections. Another consideration accounted for overhangs and hollow sections. Specifically in the material extrusion process, there exists a “critical angle” θ in which any overhang that is less than this angle (relative to the build platform) requires material below it.

Specifically in the material extrusion process, there exists a “critical angle” θ in which any overhang that is less than this angle (relative to the build platform) requires material below it. When it hangs over a hollow section, support structure is required. This was encountered in two regions with the desired print orientation, and care was taken in the design of those regions to exceed the critical angle (Figure 3). With this design and print orientation, no support structure was needed for the prototype and the entire heat exchanger was able to be printed as a single piece, thus eliminating any excess material or post-processing.Build directionFigure 2: Representation of print orientation.Figure 3: Critical angle illustration.2253One problem that was encountered, however, was the orientation in which the fins inside the air channels were printed.

The first prototype was designed with arrays of 1-mm round pin fins, but it was suspected that they would not be able to be printed perfectly round. A computed tomography (CT) scan was done on the first prototype to investigate the shape of the fins as printed. Figure 4 shows a scan in the cross-section of a fin array. It can be seen that not only are the fins not round, but they are irregularly shaped, and inconsistencies exist. The average diameter of each fin is approximately 1 millimeter, as designed, but the surface area would be greater than a round fin due to the apparently bumpy surface.Further analysis is planned for evaluation of these irregularly-shaped fins. The implications of these inconsistent fin shapes are unknown, and in order to better predict how other fin shapes and sizes would be printed, a test piece with a variety of fin shapes has been printed and scanned for a printability study.

These shapes vary from round to elliptic to airfoil, and they also vary in sizes ranging from 0.4-mm to 2.4-mm in the flow direction. Analysis of these shapes will help determine any trends in which some shapes are printed, potential design techniques to combat irregularities, and possibly a certain minimum feature size that is deemed unattainable with the current methods of printing.1 mmFigure 4: Cross-sectional view of pin fin array in CT scan.2254Analytical Performance SimulationAn analytical model was developed using Engineering Equation Solver (EES) software (Klein, 2015). The purpose of this model is to evaluate the performance of the cross-flow heat exchanger shown in Figure 1.Some assumptions were made to develop this model.

The purpose of this model is to evaluate the performance of the cross-flow heat exchanger shown in Figure 1.Some assumptions were made to develop this model. Because the temperature and pressure changes expected for both fluids are relatively small, their thermodynamic and transport properties do not change significantly within the heat exchanger; therefore, these properties are treated as being constant and are evaluated at each fluid’s average temperature. The heat capacity rate of the water will be significantly greater than that of the air and therefore the temperature drop of the water is small compared to the temperature rise of the air; therefore, the surface temperature inside the air-side channels is treated as spatially uniform. In a steam condenser the temperature will be nearly constant on the water side as long as pressure drop is small.

In a steam condenser the temperature will be nearly constant on the water side as long as pressure drop is small. The apparatus used to provide flow to both sides of the heat exchanger must be designed to uniformly distribute the flow; it is therefore assumed that there are no flow non-uniformities in the heat exchanger. The exterior of the heat exchanger will be well-insulated so that the rate of heat loss to the environment is negligible; the model assumes no heat transfer between the heat exchanger and its surroundings.The heat transfer rate of the heat exchanger and the air-side pressure drop are among the parameters of most interest. The heat transfer rate is calculated with the epsilon-NTU heat exchanger analysis method using the conductance of the heat exchanger as calculated by a thermal resistance network. The network incorporates the convection on the water-side channels, the conduction through the walls, and the convection on the air-side channels.

The network incorporates the convection on the water-side channels, the conduction through the walls, and the convection on the air-side channels. The convection on the air-side channels includes both the channel surfaces and the extended surfaces of the pin fins.One difficulty that is encountered is estimating the heat transfer coefficient and pressure drop within the air-side channels specifically. Though the flow over the pin fins can be approximated using known external flow correlations for a bank of tubes (Žukauskas, 1972) it is not an exact representation of the actual flow conditions because of the small size of pin fins and the presence of the surrounding walls. To achieve a better estimate of the heat transfer coefficient and pressure drop for this geometry, computational fluid dynamics (CFD) will be used.

To achieve a better estimate of the heat transfer coefficient and pressure drop for this geometry, computational fluid dynamics (CFD) will be used. The results from the CFD simulations will then be embedded within the analytical model in EES.Experimental PerformanceA prototype heat exchanger was printed using unfilled ABS material (Figure 5) to demonstrate the printability of the heat exchanger and to validate the model. The prototype was designed with eleven air channels and six two-pass water channels. An array of 1-mm pin fins were printed inside the air channels and all walls were designed with 0.8-mm thickness. The thermal conductivity of the material was approximately 0.2 W/m-K. The prototype was tested over the velocity range shown in the table and the heat transfer results were compared with the analytical model.

The prototype was tested over the velocity range shown in the table and the heat transfer results were compared with the analytical model. The experimental heat transfer rate is calculated using energy balances on both the air and water side.2255Table 1: Experimental operating conditionsPressure (both sides)AtmosphericInlet Temperature Air 22°CInlet Temperature Water 60°CInlet Velocity AirFlow Rate Water0.35 – 1.35 m/s 0.35 LPMMaterial Conductivity0.2 W/m-KFigure 5: Prototype heat exchanger printed using material extrusion with ABS.The experimental data and the model prediction of that data from the model correlates well for the range of air velocities tested. This demonstrates an agreement between the model and experimental results for a low-conductivity case.

This demonstrates an agreement between the model and experimental results for a low-conductivity case. The pressure drop was also measured and compared to that predicted from the model, which shows that the calculated and experimental pressure drops deviate by a factor of about 20% at the higher velocities.

This deviation is likely caused by variations between the model and the print itself, as previously described in the “Printability Considerations” section, specifically referring to the size and shape of the pin fins.2256)W( e t a rr e f s n a r tt a e H) a P ( p o r d e r u s s e r P9085807570656055504540706050403020100AnalyticalExperimental (water) 1Experimental (water) 20.30.40.50.60.70.80.9Air velocity (m/s)Figure 6: Experimental thermal data.AnalyticalExperimental 1Experimental 20.30.40.50.60.70.80.9Air velocity (m/s)Figure 7: Experimental air-side pressure data.225711OptimizationThe first heat exchanger design used an array of pin fins within air channels to establish a starting point for design, analysis, and manufacturing. The project also incorporates a method of topology optimization.

The project also incorporates a method of topology optimization. Traditional topology optimization is typically used for structural applications, but it can also be used for thermal-fluid type of work. Its most basic principle involves dividing a certain design space into unit cells, similar to a fine mesh. Then, with a code in which the simulation is performed, each unit cell is deemed to be either solid or empty. This analysis is initiated with an initial geometry and boundary conditions similar to CFD. The application with the air-cooled heat exchanger is to model the air channel as the design space and maximize the heat transfer rate while minimizing pressure drop. The goal would be for the topology optimization to produce potentially non-intuitive fin geometries, which might not be practical to manufacture using conventional methods, but could be manufactured using material extrusion.

Such geometries are expected to be presented in future work.ConclusionsThe parallel efforts between manufacturing, design, analysis, and optimization groups is crucially collaborative since motivations by one group can be limited by constraints from another; this ongoing work is a continually iterative process.Limitations on the performance of 3D printed heat exchangers go beyond the thermal conductivity of the material with which they are printed. The effort to enhance the printability of small features and thin walls should be conducted alongside that of improving the thermal conductivity.The model for analysis of the heat exchanger has been partially validated by experimental results. Upcoming designs and tests will be done to further observe and improve the accuracy of predicted performance.ReferencesJoen, C.T., Park, Y., Wang, Q., Sommers, A., Han, X., & Jacobi, A. (2009). A review on polymerheat exchangers for HVAC&R applications, Int. J.

Žukauskas, A. (1972). Heat transfer from tubes in crossflow. Advances in Heat Transfer, Vol. 8.J.P. Harnett & T.F. Invine, Jr. (Eds.). New York, NY: Academic Press.2258AcknowledgementThis work is supported by the ARPA-E ARID program with the collaboration of Oak Ridge National Lab. The support of Patrick Hruska, Tom Mulholland, and Professor Sanford Klein also made this work possible through their support at UW-Madison.2259.

sign and Use of a Penetrating Deposition Nozzle for Z-Pinning Additive ManufacturingBrenin Bales1, Roo Walker1,2, Deepak Pokkalla1,2, Seokpum Kim2, Vlastimil Kunc2, Chad Duty1,21University of Tennessee Knoxville, TN2Manufacturing Science Division Oak Ridge National Laboratory Oak Ridge, TNAbstractFused Filament Fabrication (FFF) involves depositing material layer-by-layer to create athree-dimensional object. This method often demonstrates high mechanical anisotropy in the printed structure, leading to a drop in the material strength of the part when comparing structures along the deposition plane (X/Y-Axis) versus across layers in the build direction (Z-Axis). Initial efforts to improve anisotropy led to the development of the Z-Pinning process, where continuous pins are deposited across layers in the Z-Axis. Z-pinning has demonstrated significant gains in toughness and inter-layer strength, particularly in fiber-reinforced materials.

Z-pinning has demonstrated significant gains in toughness and inter-layer strength, particularly in fiber-reinforced materials. However, this process can also create flaws in the structure that increase in severity and frequency as the pins grow in length and diameter. To mitigate this, a penetrating nozzle has been developed that extends a fine-tipped extrusion nozzle deep into the pin cavity and simultaneously extrudes material as it retracts. This study investigates the printability of the penetrating nozzle for simple geometries and evaluates the resulting Z-pinning mesostructure. As a result of this study, the prototype penetrating nozzle design was analyzed and built. Through a pressure driven flow analysis it was determined that filament will flow through the penetrating nozzle as the system pressure drop of 9.3 Mpa is less then the minimum critical pressure of 12.07 Mpa.

Additionally, it was after a transient thermal simulation, it was found that after a pause of 15 seconds the system can resume printing with no drop in heat at nozzle exit. This means the additional length of the penetrating nozzle, will not cause any clogs during any pauses in filament flow.This manuscript has been authored in part by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes.

The Department of Energy will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).1163IntroductionAdditive manufacturing (AM) is one of the fastest growing technologies for both metalsand polymers in the manufacturing industry. Primarily due its lack of tooling required, design freedom, and lower cost per part created [1]. One of the most popular methods of additive manufacturing is the material extrusion process known as fused filament fabrication, first pioneered by Scott and Lisa Crump as fused deposition modeling (FDM) with Stratasys [2]. The FFF/FDM method involves extruding thermoplastic (in a filament feedstock) through a nozzle to create a printed structure build along the X and Y axis on a build platform. The material extrusion occurs in a layer-by-layer fashion on the build plate, building up in the Z-axis.

The material extrusion occurs in a layer-by-layer fashion on the build plate, building up in the Z-axis. Due to the layer-by-layer build pattern, FFF printed structures are highly anisotropic, leading to drops in strength. This drop in strength is most severe when the load direction is orthogonal (90⁰) to the print direction (Z-axis) The best and strongest material properties are seen then the printed structures are tested so that their material beads are aligned (0⁰) in the load direction[3–5]. Due to the high anisotropy seen when comparing samples in the Z-axis (across build layers) to X-Y plane, forces are transmitted through the adjoining layers in the Z-axis along discrete interfaces between material beads on the X-Y plane [6]. Depending on the printing system and material the reduction in strength can range from 25% to 97% [7–13].

Depending on the printing system and material the reduction in strength can range from 25% to 97% [7–13]. Specifically, across both large (BAAM and LSAM) and small scale systems PLA shows a drop in strength ranging from 33% to 57% across the X and Z-axis [7,8]. Neat ABS and Fiber Reinforced ABS across BAAM and LSAM large scale systems and small scale FFF systems shows a greater range of strength reduction. Ranging from 25% to 97% reduction in ultimate tensile strength [7,9–13].Researchers at a variety of institutions have developed methods to combat the highanisotropy found in FFF printed structures. Primarily the methods attempt to increase the inter- layer bond strength by increasing inter-diffusion of molecular chains across layers. Increasing the inter-diffusion of polymer chains across the interface can be controlled by keeping the print temperature of the material above the glass transition temperature of the material [14–16].

An example can be seen in a collaboration between the University of New Mexico and Texas- Austin. The method developed by the two university research teams demonstrated a 50% increase in ultimate tensile strength across print layers (Z-axis) by increasing the extrusion multiplier values over their stock settings. Additionally, this method saw a 10% increase in X- axis ultimate tensile strength and the ability to print near isotropic structures [17]. Other research has been focused on adding a pre-heating method just before deposition of new material. Oak Ridge National Laboratory (ORNL) developed a process for the BAAM (Big Area Additive Manufacturing) system in which infrared laps are mounted to the printing head. This system led to a doubling of the fracture energy in 20% CF-ABS printed structures [18].

This system led to a doubling of the fracture energy in 20% CF-ABS printed structures [18]. A research group from Arizona State University developed a method to increase the temperature of the build surface by using a laser aimed at the surface preceding where the filament will fall. This method increased the flexural strength of 0.03% Carbon Black ABS by 95% [19]. A team from the University of Tennessee and ORNL sough to alleviate the high anisotropy and drop in strength through the creation of the novel Z-pinning method, which was introduced at the 2017 SFF conference [20].Z-Pinning1164The Z-pinning for AM process can be largely broken down into four steps that define how the structure is printed. The AM Z-pinning process will start with the deposition of the surrounding structure (the structure that is being pinned). For testing purposes this surrounding structure is a rectilinear grid, as seen in Figure 1a.

For testing purposes this surrounding structure is a rectilinear grid, as seen in Figure 1a. During this stage, a predetermined number of layers are printed (set by the pin length) and the pin cavities are created by leaving gaps in the structure itself. Figure 1b demonstrates the pinning process, here the nozzle will be positioned directly over the pin cavity and deposit material to fill a prescribed amount of the pin cavity. This process is then repeated in Figure 1c and 1d during these stages more of the surrounding structure is printed, and another pin is deposited at the alternative position. This process repeats for the entire structure (Figure 1e and 1f).

This process repeats for the entire structure (Figure 1e and 1f). It should be noted that the pins are staggered so the seams of the pins are not all on one layer.Figure 1: The AM Z-Pinning Process [21]Z-Pinning ParametersThe pin dimensions that determine how much existing structure is deposited or howmuch material is extruded into the pin cavity make up the Z-pinning parameters used in the Z- pinning process. The Z-pinning parameters are defined as the width and length of the pin, as well as the infill percentage and seam offset (See Figure 2). The pin widths used in AM Z-pinning are defined as an integer multiple of the exit diameter of the nozzle, which is defined as “W” (W=0.5mm). For the initial Z-pinning studies the typical tested diameters were 1.5mm, 2mm, and 3mm. These are defined in studies as 3W, 4W, and 6W respectively. The second pin parameter that is defined is the pin length.

These are defined in studies as 3W, 4W, and 6W respectively. The second pin parameter that is defined is the pin length. This is simply the number of layers that the pin will occupy (Figure 1b and 1d for example). Initial investigations were focused on three pin lengths: an 8-layer, 12-layer, and 20-layer pin. These are defined in combination with the seam offset that is applied so that all the pin interfaces do not form one line across the structure (See the circles in Figure 2). The seam offset for all studies is held at half the length of the pin (L). This means1165that pins are deposited in the other pin position after L/2 layers of structure have been printed. The final pin parameter is the infill percentage of the pins themselves. This is the amount of material that is extruded in the pin cavity. It is defined as the amount of material deposited compared to the theoretical volume of the cavity in which it is being deposited.

This is the amount of material that is extruded in the pin cavity. It is defined as the amount of material deposited compared to the theoretical volume of the cavity in which it is being deposited. For the AM Z- pinning process three infill percentages were investigated, 80%, 100%, and 120%. Overfill at 120% was possible due to the sparse nature of the rectilinear grid that is printed as the surrounding structure (See Fig. 2).Figure 2: AM Z-pinning parameters [22]Initial Studies in Z-PinningInitial work in Z-pinning for additive manufacturing focused on reducing anisotropyfound in FFF printed structures by improving the Z-axis ultimate tensile strength, leading to the development of quasi-isotropic properties [21]. Of the pin parameters studied, the pin length, width, and seam offset were kept constant with only pin infill increasing from 0%-120%. Additionally, a solid sample was also produced to determine maximum strength of the sample [21].

Additionally, a solid sample was also produced to determine maximum strength of the sample [21]. This study was able to demonstrate that as the pin infill increased from 0% to 120%, toughness, stiffness, and ultimate tensile strength all increased. The data also showed that as the Z-axis ultimate strength increased with pin infill, the degree of anisotropy present in the structure would decrease as well [21].Subsequently, another round of research was conducted that expanded upon the initial pinparameter set. This work investigated larger pin diameter of 2 mm and 3 mm, while also increasing the pin lengths tested from 8 layers to 12 and 20 layer pins [23]. This initial study further reinforced initial findings: as pin infill increased, so did the mechanical properties in both X- and Z- directions. Additionally, it was discovered that increasing the pin diameter and length1166increased tensile properties as well as lowered the degree of anisotropy present in the samples [23].

Additionally, it was discovered that increasing the pin diameter and length1166increased tensile properties as well as lowered the degree of anisotropy present in the samples [23]. This can be seen most prevalent in the 80% sample set shown in Figure 3. Shown above in Figure 3, this 80% pin infill samples with a pin diameter of 1.5mm (3W) showed an increase of strength across Z-axis when moving from an 8 layer pin to a 20 layer pin. This again could be seen in the 2mm diameter pins (4W) in the X-axis ultimate tensile strength [23]. Yet, one of the chief results of this study was the discovery and further reinforcement of the fact that defect propagation negatively impacted the effectiveness of longer and wider pins.Figure 3: Ultimate Tensile Strength of 80% Infill Samples [23]DefectsResults from the initial rounds of AM Z-pinning concluded that as pin infill,length, and width increased, as did the mechanical performance of the print.

Yet, defects in the printing process saw a decrease in mechanical performance where greater performance was expected (See Figure 4). To understand the defects, it is important to separate the Z-pinning process into two phases. The first phase is the rectilinear grid, in which nearly no defects occur. The second phase is the pin deposition. The pin deposition is where the error propagation occurs in the AM Z-pinning process. For a successful pin deposition to occur, material must be dropped cleanly and completely into the pin cavity (See Fig. 2). Shown in Figure 5 are the two flaws during the AM Z-pinning process. First is the underflow condition (See Fig 5a), otherwise known as “rope in a bucket” [24]. This process occurs when material does not flow cleanly into the pin cavity, this causes small voids that span the entirety of the pin cavity (see Fig. 4). The second defect (Fig 5b) is the over-fill condition, otherwise known as “short shot”.

4). The second defect (Fig 5b) is the over-fill condition, otherwise known as “short shot”. This error occurs when the material does not fall completely into the pin cavity. This leads to a large defect at the bottom of the pin, while simultaneously leading to a large amount of material bonding on the print surface. In addition to the drops in material strength that occurs due to the voids that are1167created in pin cavities, the short shot error will cause an excess of material to build up on the print surface. This will lead to a total print failure when two much material will harden and cause the nozzle to run into the nodule and knock the print over.Figure 4: Voids in Z-pinning process [25]Figure 5: Flaws during the printing process: Creation of voids (A) and Overflow (B) [24]Penetrating Nozzle ApproachAs a result of the defects found in the AM Z-pinning process a modification to the FFFextruder hot end to alleviate the two failure modes.

This solution was the novel penetrating nozzle [24]. The penetrating nozzle approach would involve developing an extended nozzle to penetrate the pin cavity and simultaneously deposit material while retracting (See Figure 6). The goal of this novel approach would be to create longer Z-pins that penetrate more completely into deeper pin cavities. Having a more complete fill of the pin cavity would in theory allow for a higher successful infill percentage as well as greater mechanical interlocking with the walls of the cavity [26].1168Figure 6: Penetrating Nozzle Approach [24]When making such significant modifications to critical component of the extruder, it is important to understand the effects the extended nozzle design might have on the hotend. Two questions had to be answered during the design process to inform how the nozzle should be built and could function.

Two questions had to be answered during the design process to inform how the nozzle should be built and could function. First, can material flow through the nozzle? And second, can the nozzle and hotend stay hot during different stages of the printing process, especially during a pause in printing?Design ProcessTo answer these questions, two experiments were conducted. First, a pressure drop analysis predicted whether liquid filament would be able to flow through the nozzle. This involved measuring the rheological properties of the polymer and evaluating a viscoelastic printability model at various extrusion conditions [27]. Then, a transient thermal analysis was conducted using a 3D model of the penetrating nozzle design. These results were steady state thermal analysis and by extruding material through a penetrating nozzle.Before the pressure drop analysis or thermal analysis could be conducted, a prototype ofthe nozzle was designed.

The printer platform selected for this project was a Prusa Original MK3s+. This printer allows for a high degree of repeatability while still allowing easy customization. The penetrating nozzle design was composed of two discrete additions to the hotend. First, the nozzle selected was a 1/8 NPT 22 Gauge Stainless Steel Dispensing Needle with a nickel-plated brass base [28]. This nozzle was chosen not only for its similar exit diameter to stock FFF nozzles (0.4mm), but also the combination of materials. Nickel-plated brass in the threaded base would allow for improved thermal performance where the softer material would not be subjected to high stresses. The stainless-steel nozzle would benefit from higher material strength, which is critical considering how thin the walls of the nozzle are. The second addition to the nozzle was a 1/8 NPT to M6 threaded adaptor [29] to interface with the stock heater block and provide the beneficial thermal properties of Nickel-plated brass.

The second addition to the nozzle was a 1/8 NPT to M6 threaded adaptor [29] to interface with the stock heater block and provide the beneficial thermal properties of Nickel-plated brass. Seen in Figure 7, the penetrating nozzle was then cut down to a final length of 5mm, allowing for 20 layers of full penetration. This allows the penetrating nozzle to penetrate up to the longest pins previously studied and allows longer pins to be studied.1169Figure 7 Penetrating Nozzle PrototypePressure Driven Flow AnalysisThe pressure driven flow analysis was based on a printability model that incorporates a variety of material extrusion parameters in order to create a model of how “well” a printer can extrude material [27]. This occurs by defining the three critical pressures that the system will experience and determining the minimum of the three critical pressure values.

This occurs by defining the three critical pressures that the system will experience and determining the minimum of the three critical pressure values. This will be the limiting force that can be applied during material extrusion in order for material to flow successfully. The three critical pressures used in this experiment are motor driving, motor drive with slip, and finally Euler’s buckling pressure.Critical Pressure ValuesThe first step in the pressure driven flow analysis is to determine the three critical pressure value that the system will experience. To do this a variety of parameters from the motor and extruder design need to be determined. The equations described below refer to the three pressures that the system will experience. Motor driving pressure without slip is the maximum torque of the motor (𝐹(cid:3031)) divided by radius of the driving wheel (𝐴(cid:3033)).

Motor driving pressure without slip is the maximum torque of the motor (𝐹(cid:3031)) divided by radius of the driving wheel (𝐴(cid:3033)). If this force exceeds the sheer strength of the filament being extruded, a second force must be investigated. This is the motor driving pressure with slip, where the maximum shearing force must be calculated (𝐹(cid:3033)).1170The last force that is investigated is the buckling pressure. If the buckling pressure is the minimum value, then the filament material will buckle under the force during the extrusion process [27,30]. The equations and values for these can be seen below in Table 2. As can be seen in Table 2 the minimum critical pressure is motor driving with slip, meaning that as soon as the motor pushes over 12.07 MPa onto the filament, the drive gears will “slip” causing material to fail to flow.

Once this minimum critical pressure was determined the next step was determining the maximum pressure drop that the custom penetrating nozzle system will experience.Table 1: Critical Pressure Values [27,30]Critical PressuresMotor Driving Pressure (no slip)𝐹(cid:3031) 𝐴(cid:3033)18.88 [Mpa]Motor Driving Pressure (slip)𝐹(cid:3033) 𝐴(cid:3033)12.07 [Mpa]Euler’s Buckling Pressure𝜋(cid:2870)𝐸𝑑(cid:3033) 16𝐿(cid:2870)461.78 [Mpa]Pressure Drop AnalysisTo better understand the printability and rheological behavior of neat PLA within this novel nozzle design, the material was analyzed using parallel plate rheology on a Discovery Hybrid Rheometer (DHR). An oscillatory frequency sweep at 225 C from 0.1-628 rad/s was performed on a 25 mm plate with a 1.0 mm gap and a strain percentage of 0.5% [30]. It was found that the neat PLA was shear thinning above 100 rad/s, which determined the region of interest when analyzing the pressure drop.

It was found that the neat PLA was shear thinning above 100 rad/s, which determined the region of interest when analyzing the pressure drop. The pressure drop equation is defined below in [1]. To calculate this, the viscosity and shear rate through the nozzle must be determined.∆𝑃 =8𝜂𝑄𝐿 𝜋𝑅(cid:2872)Most additive manufacturing polymers are shear thinning which implies that the shear rateinfluences the viscosity (η). This can be described with a power law relationship defined by:𝜂 = 𝐶𝛾̇ (cid:3041)(cid:2879)(cid:2869)where n is the power law index and C is a constant called the co. The Rabinowitsch-corrected shear rate through a nozzle [32] is defined by:1171[1][2]γ̇ =4𝑄 𝜋𝑅(cid:2871) (cid:3436)3𝑛 + 1 4n(cid:3440)Utilizing the above power law related equations [2] and [3], the pressure drop [1] can be found.

The pressure drop equation [1] and the power law viscosity equation [2] utilize steady-state viscosity (η), in order to relate the complex viscosity (η*) values that were observed with parallel plate rheology to the viscosity experienced during FFF processing, the Cox-Merz rule was applied [Eqn. 4]. The Cox-Merz rule states that the shear-rate dependency of steady-state viscosity (η) is equivalent to the angular frequency (ω) dependency of the complex viscosity (η*) [33].𝜂(𝛾̇) = 𝜂∗(𝜔)(cid:3082)̇ (cid:2880)(cid:3104)The shear rate and viscosity values will change based on their location in the extruder,shown in Figure 8. These values are gained through inputting the data gained from rheological tests into the series of equations listed above (Equations [1-4]). The results of these are shown in Table 3. These values are then inputted all into the pressure drop equation [1]. This is the final step in the pressure drop analysis and yielded a value of 9.3 MPa.

The results of these are shown in Table 3. These values are then inputted all into the pressure drop equation [1]. This is the final step in the pressure drop analysis and yielded a value of 9.3 MPa. As a result of our system pressure drop being less than the critical pressure derived in Table 2, the neat PLA filament can flow through the system.Figure 8: Section view of Penetrating Nozzle1172[3][4]Table 2: Rheological Parameters used in CalculationsParameterEntry (A)Middle (B)Exit (C)C [PaSn]19,994Shear Rate [1/s]0.8725.9808.4Viscosity [PaS]2082730252Transient Thermal AnalysisThe next step was modeling the polymer flow through the extruder during the Z-pinning process investigate how the process might affect filament heating/cooling and flow. In order to do this a model of the hotend was created on AutoDesk CFD 2021 with a volumetric flow rate that mimics the printing and pause steps that occur during the Z-pinning process.

In order to do this a model of the hotend was created on AutoDesk CFD 2021 with a volumetric flow rate that mimics the printing and pause steps that occur during the Z-pinning process. The print – pause – print transition steps are critical to understand because of the extra length of the nozzle will affect how thermal characteristics of the polymer (due to travelling farther from the heating element). In this step, the nozzle will have just finished fresh pin layer before moving back across the surface of the structure slowly back to the home position at the start of the printed structure before the deposition of another layer will occur. This step will roughly take 15s in which no filament is being extruded, but the heating element still holds the printing temperature.

This step will roughly take 15s in which no filament is being extruded, but the heating element still holds the printing temperature. The purpose of this study is to investigate the thermal properties at the very tip of the nozzle, in order to ensure that the filament at the nozzle tip will not cool below 180⁰C, which was determined to be the point in which filament flow is inconsistent and material will not extrude on the stock nozzle.Model Creation and SetupThe model created for the simulations was the prototype hot end assembled on the PrusaFFF system (See Figure 9). The materials assignments in the 3D model set according to the prototype’s construction and properties were defined by the standard values found in Autodesk material library. This was then experimentally verified by conducting a steady state analysis with no filament present meaning no material flow, just a static heat value and boundary conditions.

This was then experimentally verified by conducting a steady state analysis with no filament present meaning no material flow, just a static heat value and boundary conditions. This was verified by creating Arduino thermistor circuit and measuring various points around the base of the nozzle and adaptor, they were found to be within 2⁰C of the steady state simulation [34]. Shown in Table 4, the materials chosen for the project were 304-Stainless steel for the penetrating nozzle and heartbreak, brass for the threaded adaptor and nozzle base, and standard aluminum for the heater block.

Additionally, due to the rheological analysis performed for the pressure drop analysis, the filament was fully and accurately modeled.1173Figure 9: Prototype and CFD Model and Material LocationsTable 3: Material Assignments for Prototype NozzleMaterial 304-Stainless SteelAssignment Nozzle (A) and Heatbreak (B)BrassThreaded Adaptor (C) and Nozzle Base (D)AluminumHeaterblock (E)Simulation MethodologyThe model was constrained in order to accurately model the effect that not only the surrounding air and environment will have on the heat end, but what the filament will do as it is extruding during the printing process. As such, both radiation, film coefficient, and convection are modeled. Radiation emissivity values were applied to all surfaces (external and internal faces) according to the standard values of the material present in combination with the background temperature of the face.

The film coefficient was applied to all faces in which contact with air was possible and set according to the standard values of the various materials, as well as the background temperature. As a feature of Autodesk CFD’s solver, conduction was applied for all materials in the build. The only external surfaces that did not have a radiation or a film coefficient was the heaterblock (see Figure above). This was done to keep the temperature of the heaterblock itself at a consistent 225⁰C (the printing temperature). To do this, a static1174temperature value was applied to all external and internal faces of the heaterblock. In addition, the exit and entrance temperatures of the heatbreak were set mimic how the what temperature gradients the filament will experience as it is traveling into the liquefier zone inside the heaterblock [35]. This ensures that as it enters and exits the liquefier zone the filament will maintain the temperatures it does during the printing process.

This ensures that as it enters and exits the liquefier zone the filament will maintain the temperatures it does during the printing process. Convection was not applied to the hotend due to the hotend fan not running during Z-pinning, as a result there was no forced convection boundary conditions present. Lastly, a volumetric flow rate is set to allow the filament to pause while the extruder is heating, print for 270 seconds, pause for 15 seconds, then finally printing again for 120 seconds, a graph of the volumetric flow behaviors is shown below in Figure 9. The flow rate itself is set for the 4.792 mm3/s, which is the lowest flow rate the material is deposited at during the process. A full model, with constraints can be seen below in Figure 10.Figure 9: Volumetric Flow Rate Values, Units in mm^3/sFigure 10: Full Penetrating Prototype, with constraints1175The model was meshed by Autodesk CFD using the automatic meshing software tetrahedralelements.

The maximum size for all edge elements created was 10 mm, with a minimum refinement length of 0.037 mm. While surface meshes had a maximum size of 0.002 mm. Then an automatic surface refinement was applied to all surfaces to reduce the element count to 86k elements (both solid and fluid). Once the model was meshed , the simulation was run as a transient solution with a step size of 0.08 seconds with 3 inner iterations per time step. This resulted in a full convergence of all values calculated by the solver.ResultsAfter running the simulation for the full time span the model converges. Figures 11 and 12 are the thermal maps of the penetrating deposition nozzle and extruder. Shown in Figure 11 is the extruder right at the end of the printing process, specifically one second away from the pause stage (See Fig. 9).

Shown in Figure 11 is the extruder right at the end of the printing process, specifically one second away from the pause stage (See Fig. 9). Clearly shown on the left end on the heatbreak, the filament is still solid while at 35⁰C, yet as it enters the heaterblock, it will rapidly heat up and become liquid at 225⁰C. Figure 12 critically, shows the thermal map of the extruder after the pause stage ends. It is important to note that due to fresh, cold filament not being pushed through the extruder, the stationary filament is able to entirely heat up to 225⁰C. Due to the lack of extrusion this heat will then creep back up the heatbreak and into the heatsink (not shown).

Due to the lack of extrusion this heat will then creep back up the heatbreak and into the heatsink (not shown). In a full experiment, a fan will be positioned perpendicular to the heatsink in order to keep the entire heat sink (and filament inside it) to roughly 30⁰C - 40⁰C [35].Figure 9: Thermal State of Prototype Extruder (T=299s)1176Figure 10: Thermal State of Prototype Extruder (T=316s)It can be clearly seen in Figures 9 and 10 that no matter how long the print is paused, due to the lack of cold filament being driven in to the hotend, the entire assembly will continue to heat up. In fact, if there was no heatsink or fan attached to the extruder, there would have been negative effects due to the filament melting to high in the hotend/heatsink. A graphical representation of this can be seen below in Figure 11.

A graphical representation of this can be seen below in Figure 11. Specifically, this graph looks at the exit temperature of the nozzle to determine if the filament ever cools down to below 180⁰C during the pause stage of the printing process. As you can see in the graph below, there is no point in which the filament will cool below to below 200⁰C, let alone 180⁰C. Important to note is that the end of the nozzle probe conducts pure filament and the nozzle, so the rapid ramp up in the first 100 seconds is due to no filament actually reaching the point, and the average temperature being calculated against the filament and wall of the nozzle.Figure 11: End of Nozzle Temperatures during Simulation1177Conclusions and Future WorkThe Z-pinning process for additive manufacturing has already shown in initial work to be an effective method for increasing Z-axis ultimate tensile strength as well as reducing anisotropy found in printed parts.

However, defect propagation stymied the ability of longer and wider pins to effectively deposit good pins, reducing mechanical performance where increases were expected. As a result, a penetrating nozzle approach was proposed. The penetrating nozzle approach would combine the standard material deposition of stock FFF nozzles, while also introducing the ability to penetrate the pin cavity in order to lay a cleaner Z-pin deeper into the structure.In order to quality a design before test prints could be conducted, two questions had to beanswered before serious prototyping occurred. First, can material flow through the nozzle? Second, if material can flow, can it pause and then resume? The first question was answered by using a pressure drop analysis pioneered to quality how “printable” a system was. As a result of the rheological and geometric analysis of a mock-up of the penetrating nozzle, it was discovered that the critical pressure of the system was over 12Mpa.

As a result of the rheological and geometric analysis of a mock-up of the penetrating nozzle, it was discovered that the critical pressure of the system was over 12Mpa. Which, when compared to the system pressure drop of 9.338Mpa, means that filament will flow through the system. This was verified through loading the material and performing a simple extrusion command in G-Code. To answer the second and final question, a full CFD analysis of the filament flow through the nozzle was conducted. As a result of this analysis, it was determined that once the extruder has been printing, the pause that occurs in standard Z-pinning is not long enough for significant cooling to occur. In fact, it was discovered that due to the nozzles proximity to the heating element, it will in fact continue to heat up until it reaches the system temperature of 225⁰C. This was experimentally verified as well by pausing the printer after loading material and seeing filament continue to “ooze” out.

This was experimentally verified as well by pausing the printer after loading material and seeing filament continue to “ooze” out. Therefore, showing that even when the extruder had stopped depositing filament, the nozzle remained hot enough for residual filament to remain liquid inside the nozzle. After the pause, the extruder was able to resume deposition as well.Due to these two pieces of analysis, it can be concluded that filament will flow through thepenetrating nozzle design. For the future work in this method, microstructure analysis will be conducted to validate the effectiveness of the penetrating nozzle in regard to defect reduction as well as comparing the penetrating Z-pins, to standard Z-pins.

Mechanical tensile testing will be conducted to determine if the penetrating nozzle was able to increase the tensile properties of the samples, either through defect elimination or improved interlocking.1178AcknowledgementsSpecial thanks to Nathan Black and Alex Defilippis as well as the University of Tennessee Maker Space and Innovation and Collaboration study for their help in this project. Thanks to the U.S. Army Combat Capabilities Development Command Aviation & Missile Center as well for sponsoring this research. Additional support was generated under the National Science Foundation Award No. 2055529.This research was sponsored by the Research sponsored by the U.S. Department of Energy,Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle, LLC.1179References[1] H. Bikas, P. Stavropoulos, G. Chryssolouris, Additive manufacturing methods and modellingapproaches: a critical review, Int. J.

SIGN AUTOMATION FOR MULTI-MATERIAL PRINTING Jonathan D. Hiller, Hod Lipson Mechanical and Aerospace Engineering Cornell University Ithaca NY, 14853, USAAbstractThe advent of multi-material freeform fabrication technologies has exponentially increased the mechanical design space available to engineers. The feature-based paradigm of traditional CAD software is insufficient to take advantage of the freedom of internal material distribution and gradients. Here we present a flexible evolutionary design algorithm for 3D multi-material structures that fully utilizes this expanded design space. The material distribution is optimized subject to high level functional constraints, or simple constraints such as maximizing stiffness per weight. The algorithm is inherently capable of shape optimization, or can simply optimize material distribution within a given geometry.

The algorithm is inherently capable of shape optimization, or can simply optimize material distribution within a given geometry. We demonstrate autonomous design of freeform shapes, 3D non-uniform structures, and 3D compliant actuators.IntroductionDespite the ubiquity of computing power available, the field of mechanical synthesis has so far resisted widespread automation. At the root of this problem is the inability of algorithms to transfer multiple competing goals efficiently into a geometry which can be manufactured practically. To accomplish this, an algorithm must incorporate some knowledge about the manufacturing process to be used, and stay within the constraints of the process. These could include factors such as fixtures for a machining process or draft and parting lines for a molding process.

These could include factors such as fixtures for a machining process or draft and parting lines for a molding process. However, the advent of freeform additive manufacturing technology allows any 3D shape and topology to be fabricated without penalty (Beaman et al., 1997), removing many complex constraints from the automated design process.In addition to the lack of geometrical constraints when designing for additive manufacturing (Hague et al., 2003), with the advent of multi-material 3D printing (Malone et al., 2004, Objet, 2009) the mechanical design space available to engineers is increasing exponentially. Traditional feature-based CAD programs are generally insufficient to create optimal freeform geometries (Mantyla et al., 1996), and are severely limited in their ability to design using multiple materials both at a macroscopic and microscopic level.

For this reason, design automation will likely play an increasingly important role in mechanical synthesis, especially for parts created using additive manufacturing processes.BackgroundThe ideal design process would require only the functional goals of the desired part to be input, and would create the optimal blueprint autonomously. Much work has been done in this field (Díaz and Lipton, 1997, Fernandes et al., 1999) regarding the topology or shape optimization of single material structures, with the goal of maximizing stiffness per weight. The most established method in this field is homogenization, as originally demonstrated by Bendsoe and Kikuchi, 1988. This iterative process varies the effective stiffness of each cell within a 2D or 3D matrix according to its strain energy, and optimizes the structure subject to constraints on total volume and minimizing strain energy.

Variations on this method have yielded results that maximize deflection for applications such as a simple gripper structure, or even utilize two materials to emulate an actuated structure (Buehler et al., 2004, Nishiwaki et al., 1998, Sigmund and Torquato, 1999).However, homogenization results are generally limited to optimizing overall deflection or force. This approach become unwieldy or intractable as complexity is added to the design problem, such as competing objectives involving multiple materials or specifying a desired deformed shape.Evolutionary algorithms have also been explored for the purposes of topological optimization. This class of algorithms maintains a population of possible solutions, which are continually mutated and recombined (crossover) to improve over time. Evolutionary algorithms have not found widespread use in topological optimization for several reasons.

Evolutionary algorithms have not found widespread use in topological optimization for several reasons. First, they are much less efficient than the homogenization method for the single objective structural optimization problems that are often addressed in literature. Secondly, the success of genetic algorithms depends on how the object is encoded (the genotype) to represent the physical object (phenotype). In early attempts at using genetic algorithms to solve topological optimization problems, every individual pixel or voxel was represented explicitly in the genotype (Jakiela et al., 2000, Kane, 1996). In addition to the challenge of making crossover and mutation non-destructive, this method scales poorly to large structures.Thus, the success of genetic algorithms in solving topological optimization problems depends on using a representation that efficiently defines sensible objects using a minimum number of parameters (Kane and Schoenauer, 1996).

Various graph structures, generative encodings, and constrained bit-wise encodings have been proposed to address these challenges, as summarized in Kicinger et al., 2005. These more efficient representations enable genetic algorithms to overcome the functional complexity limitations of homogenization, although at the expense of increased computational effort.Results and Discussion Encoding (how the algorithm thinks about the objects)Here we propose an encoding which naturally generates smooth, freeform 2D and 3D shapes at an arbitrary level of complexity with a minimal number of parameters for any number of materials. The genotype consists of a series of frequency amplitudes atharmonic multiples, and the phenotype is generated by applying the inverse discrete cosine transform (DCT) to this series. This representation was chosen for its relative simplicity of implementation and evolvability.

This representation was chosen for its relative simplicity of implementation and evolvability. By representing a three-dimensional object as a 3D matrix of frequency amplitude components, a number of advantages are realized. These include the ability to render the genotype at any resolution, complete dimensional independence, and naturally available symmetry. This comes at the expense of not being able to easily represent sharp corners and flat surfaces in the generated objects. The discrete cosine transform was chosen over other transforms due to its inherent ability to concentrate the most useful information in the lowest order frequency components.The user-defined complexity metric is worth explaining in greater detail. If desired, the DCT encoding could contain as many frequency components as voxels within the physical object. This would allow any possible configuration of voxels to be generated, but defeats the purpose of having an indirect encoding.

This would allow any possible configuration of voxels to be generated, but defeats the purpose of having an indirect encoding. As we decrease the number of parameters in the frequency matrix, the ability to create any possible object is lost. However, the encoding was chosen such that sensible, smooth objects suitable for additive manufacturing would be generated with only a few parameters. The complexity metric is defined as the maximum number of frequency components in the longest dimension. If a sample workspace was 40 voxels long, a complexity metric of 10 would result in a minimum feature size of approximately 4 voxels (40/10).Fitness (How the algorithm evaluates possible solutions)An integral part of the genetic algorithm is the fitness evaluation. At this step, the algorithm evaluates each design according to user defined criteria.

At this step, the algorithm evaluates each design according to user defined criteria. Here, we are concerned with the mechanical reaction of our structures to various conditions, calculated using the linear direct stiffness method as illustrated in Figure 1. For each experiment, certain regions of voxels were defined as fixed to ground or subject to an applied force. Each voxel has an associated stiffness and poisson’s ratio, based on the output of the inverse DCT. This defines the global stiffness matrix for a given geometry, which was then solved using the highly optimized PARDISO solver (Schenk and Gärtner, 2006, Schenk and Gärtner, 2004) library to yield the resulting displacements and internal stresses of each voxel of the structure under load.(a)(b)Figure 1: The direct stiffness method was used for physical evaluation of multi-material cantilever beams.

Shown here are the deformations resulting from a homogenous material (a) and an automatically generated multi-material distribution (b).Algorithm DetailsThe genetic algorithm used here is outlined in Figure 2. When two individuals are selected for crossover, each frequency component of the offspring is randomly chosen from either parent. The parent which contributed the greatest number of frequency components is stored as the "most similar" parent for the selection process. To mutate an individual, random variations were introduced into the frequency amplitudes of the genotype. These variations were kept small, such that any value could not change more than 10% for any given mutation. For all experiments presented here, deterministic crowding selection was used with a population of 10-25 individuals and 20% probability of mutation. In this method, each individual was paired with a random mate and crossover was performed.

In this method, each individual was paired with a random mate and crossover was performed. If the child was more fit than its most similar parent, the child replaces the parent in the population. Otherwise, the offspring is disregarded.Figure 2: Flowchart of the multi-material genetic algorithm used for design automation. The discrete cosine transform encoding allows freeform objects to be encoded with a minimal number of parameters.Shape OptimizationGiven any set of loading and constraint conditions, the algorithm is successful in creating and optimizing the stiffest shape. Here, the result of the inverse DCT was simply thresholded to determine which voxels were instantiated. The case of a cantilever beam is shown in Figure 3. Here, the population size was set at 25 and the domain used while evolving was 24x6x12 voxels.

The case of a cantilever beam is shown in Figure 3. Here, the population size was set at 25 and the domain used while evolving was 24x6x12 voxels. The algorithm was allowed to run for approximately 3500 generations, although the gains past 1200 generations were negligible.(a)(b)Figure 3: A cantilever beam was autonomously designed to maximize stiffness and minimize weight (a) using the proposed algorithm. The physical printed beam is shown in (b).Multi-material gradated material optimizationThe optimization of multiple, gradated materials presents a more interesting problem which is not addressed in previous research. In these experiments, a qualitative material model was used to generalize the ability of additive manufacturing processes which simultaneously print high stiffness and low stiffness materials. A two-material system was assumed, where both stiff and flexible materials can be placed and combined in any combination at both the macroscopic and microscopic level.

A two-material system was assumed, where both stiff and flexible materials can be placed and combined in any combination at both the macroscopic and microscopic level. In effect, the material property can vary continuously between the two extremes, as is demonstrated to be reasonable in Hiller and Lipson, 2009. The resulting material property of each location is calculated from the density output by the discrete cosine transform, according to a 4th order exponential weighting.Two different cases were chosen to illustrate the use of genetic algorithms in designing composite structures with high-level functionality. First, we considered the deflected shape of the top surface of a cantilever beam.

First, we considered the deflected shape of the top surface of a cantilever beam. Secondly, we considered an actuator that maximizes the two-dimensional deflection of a beam tip subject to applied forces in the mutually orthogonal direction to the desired deflections.Deflected beam shapesGiven a cantilever beam, a variety of desired deflected profiles were selected, departing to varying degrees from the normal third-order polynomial profile. The resulting geometries are shown in Figure 4, along with plots showing the desired deflected shape overlaid onto the actual, evolved deflected shape.The algorithm was very successful in meeting these high-level goals. The straight profile (Figure 4 a & d) was considered as a simple geometry which ideally needs only simple regions of stiff and flexible material for the ideal solution.

The straight profile (Figure 4 a & d) was considered as a simple geometry which ideally needs only simple regions of stiff and flexible material for the ideal solution. The result shows the expected distribution of flexible material near the grounded edge and stiff material composing the rest of the beam, but does even better by accounting for the zero slope boundary condition imposed by the fixed end of the cantilever beam. A discontinuous slope (sharp bend) was also considered, with similarly successful results (Figure 4b & e). Also, we defined a 4th order polynomial profile involving both positive and negative curvature that matches the fixed zero slope boundary condition of the cantilever beam (Figure 4 c & f).

Also, we defined a 4th order polynomial profile involving both positive and negative curvature that matches the fixed zero slope boundary condition of the cantilever beam (Figure 4 c & f). Remarkably, the algorithm found a solution to this problem, which is in a fundamentally different material distribution domain than the other solutions.(a)(b)(c)t n e m e c a l p s i D d e z i l a m o N10123t n e m e c a lp s i D d e z i l a m r o N10123t n e m e c a l p s i D d e z i l a m r o N0.40.61.62.63.64 -0.030.470.974 -0.030.470.974.60.030.47Normalized XNormalized XNormalized X(d)(e)(f)Figure 4: Evolved geometries (a-c) show stiff regions of material as solid red transitioning to transparent yellow for lower stiffness.

The results for the deflected shape of each evolved cantilever beam (d-f respectively) demonstrate the ability to control the deflected profile of the cantilever beam to a high degree of accuracy, including slope discontinuities and non-intuitive upward curvatures.Two-Force actuatorThe optimization of a planar actuator was also considered. The loading conditions are shown in Figure 5. Two input forces are applied midway down the beam parallel to the major axis of the beam, and the output was the displacement of the tip of the beam in the two orthogonal directions to the major axis of the beam. Fitness was defined as the reachable area of the center of the tip of the beam.0.97(a)(b)(c)Figure 5: The design automation algorithm is able to solve a high-level, non-intuitive compliant actuator problem. Given the loading conditions for actuator optimization (a), the solution achieves a large reachable area (b) from the evolved geometry shown in (c).

Given the loading conditions for actuator optimization (a), the solution achieves a large reachable area (b) from the evolved geometry shown in (c). Here, red represents stiff material and transparent yellow is flexible.In the solution, the algorithm came up a material distribution that make intuitive sense upon examination. A thin row of stiff material connects the grounded face of the beam to the mid-plane of the beam where the forces are applied. Regions of stiff material connect this strand to the two locations of applied force. The outer half of the beam essentially has no forces acting on it, so this region was effectively ignored, leading to a mostly random material distribution within it.Case Study: BracketTo further demonstrate the usefulness of this algorithm, we considered a case study where an engineer has already designed the geometry of a bracket, but would like to optimize the internal material distribution to maximize strength vs. weight.

weight. Here, we assume that the stiff material is proportionately heavier than the flexible material. The desired geometry and resulting material distribution is shown in Figure 6. The algorithm was successful, and in this case generated a material distribution that makes intuitive sense upon examination.(a)(b)Figure 6: The design automation algorithm optimizes the internal material distribution of a pre-designed bracket (a), in order to maximize stiffness and minimize weight. The results are shown in (b), where red represents stiff, dense material transitioning to transparent yellow, which represents flexible, lightweight material.ConclusionsGenetic algorithms are suitable for designing the complex multi-material objects that have recently become possible to fabricate using additive manufacturing techniques.

The lack of existing software design tools to fully take advantage of the capabilities of these fabrication processes enables genetic algorithms to fill a new niche in the mechanical design space. Instead of designing an object using traditional CAD programs, genetic algorithms allow an engineer to simply set high-level goals to be fulfilled and the blueprint is autonomously generated. Likewise, existing geometries can be optimized for multi-material printing.We have demonstrated this using several examples. The algorithm is successful at optimizing the geometry of a single material cantilever beam. Additionally, by specifying the desired deflected shape of a multi-material beam, or defining vague goals of maximizing deflection area of a compliant actuator, we solve a problem that cannot be easily addressed using current state-of-the-art tools.

Additionally, by introducing multiple objectives such as minimizing weight in addition meeting the previous criteria, we demonstrated the flexibility of genetic algorithms to easily adapt to competing objectives of very different problems. These results open the door to robust, high-level design tools for complex design problems that can fully utilize the capabilities of multi-material additive fabrication techniques.topological optimizationThis work was supported in part by a National Science Foundation Graduate Research Fellowship and NSF Creative-IT grant 0757478.References BEAMAN, J. J., MARCUS, H. L., BOURELL, D. L., BARLOW, J. W., CRAWFORD, R. H. &MCALEA, K. P. (1997) Solid Freeform Fabrication: A New Direction in Manufacturing, Kluwer Academic Publishers.BENDSOE, M. P. & KIKUCHI, N. (1988) Generating optimal topologies in structural designusing a homogenization method. Comput. Methods Appl. Mech. Eng., 71, 197-224. BUEHLER, M. J., BETTIG, B. & PARKER, G. G.

SIGN CONSIDERATIONS FOR MASK PROJECTION MICROSTEREOLITHOGRAPHY SYSTEMSPhillip M. Lambert, Earl A. Campaigne III, and Christopher B. Williams Design, Research, and Education for Additive Manufacturing Systems Laboratory Department of Mechanical Engineering Virginia Polytechnic Institute and State UniversityAbstractMask projection microstereolithography (MPµSL) uses a dynamic mask and focusing optics to digitally pattern UV light and selectively cure entire layers of photopolymer resin. These systems have been shown to be capable of creating parts with features smaller than 10µm. In this paper, the authors analyze existing MPµSL systems using functional decomposition. Within the context of a morphological matrix, these systems’ design embodiment decisions are compared and the resulting performance tradeoffs are quantified. These embodiment decisions include the dynamic mask, UV light source, projection orientation, and supporting optics.

These embodiment decisions include the dynamic mask, UV light source, projection orientation, and supporting optics. The aim of this work is to provide a design guide for the realization of future MPµSL systems.Keywords: mask projection, microstereolithography, stereolithography, vat polymerization, morphological matrix, functional decomposition, system design, system analysis, literature review1. Background and MotivationWithin the realm of Additive Manufacturing (AM), stereolithography (SL) is a well- established technology used to create parts, with features as small as 75µm, by crosslinking layers of photopolymer resin with a scanning ultraviolet (UV) laser beam. Mask projection microstereolithography (MPµSL) modifies the SL process by replacing the scanning UV laser with a UV light source and dynamic pattern generator (or dynamic mask) to digitally pattern light and expose an entire cross-sectional layer at once.

Unlike traditional SL processes, MPµSL is not limited by the laser beam radius or its scan speed, and thus enables the creation of feature sizes smaller than 10µm while also reducing build times by an order of magnitude [1–14].In 1995 Arnaud Bertsch presented the first MPµSL system, which he then called integral stereolithography [1], [2], [15] The general mask projection micro-stereolithography process flow is illustrated in Figure 1. Light is first created by a source - commonly a light emitting diode (LED), lamp, or laser. This light is then conditioned by a series of optics that may include collimating lenses, wavelength filters, and homogenizing rods. A mirror is often used to reflect light onto a dynamic pattern generator (dynamic mask), such that it is parallel to the projection surface. The dynamic pattern generator digitally patterns and projects the incident light as an image.

The dynamic pattern generator digitally patterns and projects the incident light as an image. Finally, the patterned light is resized by an optical lens, or series of lenses, to focus the final image on the surface of liquid photopolymer resin. The projected pattern initiates the crosslinking of monomers within the photopolymer resin, causing it to change phases from a liquid to a solid in a process called polymerization.Figure 1. Diagram of a typical MPµSL system. Purple lines highlight the travel of light.The first layer of photopolymer resin is cured on a build platform. The build platform is then repositioned such that additional resin recoats the previously cured photopolymer to provide material for creating a subsequent layer. An image of the next cross-sectional layer is projected to cure it on top of the previous layer.

An image of the next cross-sectional layer is projected to cure it on top of the previous layer. After the final cross-section is projected, the completed part is removed and post-processed.Many MPµSL systems have been developed, each providing the same fundamental system functionalities, but with different system embodiments to meet the requirements of a variety of applications (Section 4). This paper analyzes these past MPµSL systems and abstracts the inherent system tradeoffs in order to provide readers with a guide for designing future systems and applications. The MPµSL process is described in terms of functional components and design considerations in Section 2. Following the AM classification approach suggested by Williams, Rosen and Mistree [16], a morphological matrix design tool is used to visually categorize MPµSL design solutions. Performance tradeoffs are identified between the various system components that reoccur throughout published MPµSL systems in Section 3.

Performance tradeoffs are identified between the various system components that reoccur throughout published MPµSL systems in Section 3. The performance metrics used to quantify these tradeoffs are the systems’ (i) achievable layer thickness, (ii) minimum feature size, (iii) build volume, and (iv) vertical build time. Within the organizational context of the presented morphological matrix, this paper categorizes many MPµSL systems in Section 4, while drawing conclusions from their design decisions. The results and conclusions of this review are presented in Section 5.2. MPµSL Functional AnalysisThis section describes the MPµSL process as a set of subfunctions and performanceparameters, which are shown through a functional decomposition and a morphological matrix.2.1 MPµSL System SubfunctionsThe MPµSL process can be discretized into a set of functional subsystems, each with their own set of unique design considerations.

The final performance of a MPµSL system is dependent on the sum of these functional parts. They are as follows:Light source produces the luminous energy that is projected onto the resin surface to selectively cure photopolymer.Conditioning optics change properties of the projected light for MPµSL applications. • Dynamic mask digitally patterns and projects incident light to selectively cure photopolymer. • Projection orientation is the position of the dynamic mask relative to the polymer container. • Imaging optics expand or reduce the projected image to achieve the desired image resolution. • Recoat method supplies liquid photopolymer over previous layers for the creation of new layers. • Build platform supports the object being made. • Vertical actuator repositions the build platform for the creation of new layers. • System controls programmatically alter system properties during the fabrication process.

• Vertical actuator repositions the build platform for the creation of new layers. • System controls programmatically alter system properties during the fabrication process. • Photopolymer container holds the reservoir of photopolymer and contains the build platform. • Photopolymer is the raw material used in MPµSL to fabricate three dimensional objects.These subfunctions, and their respective design considerations, are presented as a functional decomposition in Figure 2. This list includes the core system functionalities of MPµSL systems; however, it is not necessarily comprehensive. Some applications may require application-specific considerations, such as humidity control or gas metering within an enclosed build volume to reduce the influence of environmental inconsistencies [7].This discretization of the MPµSL process into a set of necessary functional subsystems enables the categorization and comparison of MPµSL systems.

A morphological matrix – a design tool that presents system subfunctions and their respective potential design solutions – is provided for the MPµSL process in Figure 3. Using this tool, a designer is able to create different MPµSL embodiments by implementing unique combinations of subfunctions. In addition, this tool can be used as a framework for categorizing and comparing existing MPµSL systems on a functional basis.It is important to note that the presented matrix features only those subfunctions that have the most direct effect on overall system performance. The remaining subfunctions (build platform, vertical actuator, system controls, and photopolymer) are not detailed in this work as they either have only indirect effects on system performance, are common engineering components, and/or are not within the scope of this paper (e.g., controls and materials issues).Figure 2.

The MPµSL performance parameters and relationships are as follows:Layer thickness (µm) is a metric for quantifying vertical (z-axis) resolution. The achievable layer thickness is largely determined by the employed recoating method, actuator resolution, polymer characteristics, and projection orientation.Minimum feature size (µm) defines the cross-sectional resolution and smallest obtainable features in the X-Y plane. The dynamic pattern generator, projection orientation, and imaging optics define the projected cross-sectional resolution, while the recoating method restricts the physical minimum feature size.Build volume (mm3) is a metric for quantifying the dimensions of the largest printed part. The dynamic mask and imaging optics limit the size of the projected image in the X-Y plane and thus the maximum build volume.Vertical build rate (mm/s) is a metric that quantifies process throughput. Unfortunately, MPµSL build rates are often published using unclear terms.

Unfortunately, MPµSL build rates are often published using unclear terms. For example, many authors state throughput in units of seconds per layer, but do not specify whether this is the exposure time per layer or total build time per layer. This is compounded by the fact that each system uses different photopolymers, which require different exposures. This often prevents the direct comparison of vertical build rates for previously published work.A Process Planning matrix (from Quality Function Deployment methodology) is used to illustrate the interrelationships between subfunctions and performance metrics (Table 1). The strength of each relationship is indicated by assigned values where 1 represents a weak relationship, 3 a medium relationship, and 9 a strong relationship. For example: There is a strong relationship (9) between the minimum feature size (µm) and imaging optics of a MPµSLmachine, because the imaging optics expand or reduce the projected features.

For example: There is a strong relationship (9) between the minimum feature size (µm) and imaging optics of a MPµSLmachine, because the imaging optics expand or reduce the projected features. In addition, build rate is primarily determined by projection orientation and recoating method.Table 1. QFD Process Planning matrix illustrating the interrelationships between MPµSL functional subsystems and performance parameters.m u m i n i Me r u t a e F)m µ (e z i Sr e y a Ls s e n k c i h T)m µ (d l i u Be m u l o V) 3m m(l a c i t r e Vd l i u Be m Ti) s /m m(Light Source13Conditioning Optics11Pattern Light99Projection Orientation919Imaging Optics99Recoat Method9393. MPµSL System ComponentsIn this section, the authors analyze possible solutions to the MPµSL subfunctions presented in Figure 3 and Table 1.

MPµSL System ComponentsIn this section, the authors analyze possible solutions to the MPµSL subfunctions presented in Figure 3 and Table 1 . The design considerations (Figure 2) for each solution are also detailed.3.1 Light SourceThere are three performance criteria that are particularly important for selecting a lightsource for MPµSL applications:The emission spectrum of the light source must include the wavelength required by the selected photopolymer to achieve photopolymerization.The intensity of light at that wavelength must be great enough to reach the photopolymer’s critical exposure in a reasonable amount of time (e.g., < 60 seconds). This is a key factor in determining system build speed.The divergence and intensity profile of light determine the conditioning and imaging optics required to generate a focused image of homogenous intensity.Lamp-based light sources reflect light emitted in all directions from a bulb into a light guide.

These lamps output high intensity, broad-spectrum light often ranging between 350nm and 500nm. For this reason, lamp light sources are compatible with a wide variety of photopolymer resins of different wavelength sensitivities. Lamp-based light sources are the most commonly used light source for exposing photopolymers in MPµSL systems [4–9], [14], [17–24].Light Emitting Diode (LED) sources are also used in MPµSL systems [3]. In general, LED sources have longer operating lives, lower cost, smaller package size, and lower heat generation than mercury lamps and lasers. LED sources generally output lower light intensities than lamps at one or more wavelengths, which may be chosen to match the photopolymer’s polymerization wavelength. LED sources are therefore more energy efficient than mercury lamps and lasers, using energy more efficiently to photopolymerize the same volume of photopolymer [3].Lasers have also been used in MPµSL systems [1], [2], [10–13].

Laser light sources may emit light at a single wavelength or across multiple wavelengths. Lasers are available in ultraviolet, visible, and infrared wavelengths, but often cost thousands of dollars more than the other two light source options.3.2 Conditioning OpticsThe system of optical components between the light source and dynamic mask generator is referred to as the conditioning optics (Figure 1). These optical components change the properties of the original light source before reaching the dynamic mask. Unfortunately, published MPµSL systems underreport the specifics of these optical components, so it is difficult to discuss them in great detail. However, commonly mentioned components include:Homogenizing rods that internally reflect light repeatedly to create even intensity profiles. Light without a homogenous intensity profile fabricates inconsistent layer thicknesses along the projected image.

Light without a homogenous intensity profile fabricates inconsistent layer thicknesses along the projected image. These inconsistencies will compound and create dimensionally inaccurate parts [5], [24].Collimating lenses may be used to collimate a highly diverging source such as a lamp. Well collimated light diverges very little, preventing unwanted beam divergence and the resulting decrease in light density [4], [6–9], [17], [18], [23], [25].Filters may be used to remove unwanted wavelengths, isolating the desired wavelength [4], [6–9], [17], [18], [23], [25].Beam expanding optics may be used to expand already collimated light from an LED, laser, or collimating lenses [1], [2], [10–14].Different light sources produce different levels of collimation and homogeneity. Lasers produce the most collimated light with the most homogenous light intensity. Lamp light sources often possess a Gaussian intensity distribution and non-collimated, widely diverging light.

Lasers produce the most collimated light with the most homogenous light intensity. Lamp light sources often possess a Gaussian intensity distribution and non-collimated, widely diverging light. It is the authors’ experience that LED light sources are often less collimated than lasers and possess periodic drops in intensity distributions originating from their diodes.Using the optical components listed above, the collimation and homogeneity of lamp and LED light sources can be improved; however, it is difficult to achieve the same level of performance as a laser.3.3 Pattern LightA dynamic pattern generator digitally patterns and projects the conditioned light. Dynamic masks all operate by discretizing light over a 2D array of pixels, each individually controlling the light’s path. Important design considerations of the dynamic mask include its resolution, pixel pitch, and transmission.

Important design considerations of the dynamic mask include its resolution, pixel pitch, and transmission. Dynamic masks are used primarily in the digital display industry andcome in standard resolutions (e.g., 800x600, 1024x768, and 1920x1080). The pixel pitch represents the size of each pixel and the space between pixels. Because optical magnification must be uniform, the resolution and pixel pitch determines the ratio between the projected image area and the minimum feature size. This limitation can only be overcome by moving the mask in relation to the build platform, such that images are stitched together in the photopolymer [4]. Digital Micromirror Device (DMD), Liquid Crystal Display (LCD), and Liquid Crystal on Silicon (LCoS) technologies have been used as dynamic masks throughout MPµSL.3.3.1 Liquid Crystal Display (LCD)Early MPµSL systems were implemented using LCD devices as the dynamic mask [1], [2], [10–14].

LCD chips digitally pattern light by switching pixels between opaque and transparent states, achieved by controlling the orientation of the liquid crystal molecules comprising the pixel. However, the original LCD devices were not designed for use with UV light, and only transmit about 12.5% of UV light [5].3.3.2 Digital Micromirror Device (DMD)Given LCD’s limited UV transmission, the DMD was used almost exclusively in MPµSL systems after 1999 [4–10], [17–21], [23–26]. DMD’s discretize light over a 2D array of aluminum micromirrors that are individually actuated between on and off orientations (+12 degrees) by electrostatic forces applied at their hinges. The DMD offers many advantages when compared to other available dynamic masks:The DMD has small pixel sizes and narrow gaps between pixels. Because of this, DMDs are designed with greater pixel density and can reflect incident light with a more uniform intensity (less light is lost in the gaps).

Because of this, DMDs are designed with greater pixel density and can reflect incident light with a more uniform intensity (less light is lost in the gaps). Therefore DMDs have a higher filling ratio (reflective area/total area) compared to LCDs: 91% versus 57% [5].The modulation speed between states for individual pixels is also much less for DMDs as compared with LCDs: 20µs versus 20ms. This allows for greater control of exposure time and an increased ability to modulate individual pixels as to achieve gradient, grayscale projection and digitally modulate the intensity of reflected light [5].The mirrors of the DMD are surfaced with aluminum that reflects approximately 88% of the incident light, while LCDs typically transmit only 12.5% of incident light.

DMDs are therefore more efficient at patterning light and require less powerful light sources, reducing system cost and complexity [5].3.3.3 Liquid Crystal on Silicone (LCoS)Liquid Crystal on Silicone (LCoS) devices are a reflective version of LCD technologies. LCoS chips possess a 2D array of liquid crystals between one transparent thin-film transistor (TFT) and one silicon semiconductor. The transparency of each pixel is controlled by the applied voltage just as in LCD chips. Unlike LCD, the opaque pixels pass light to the underlying reflective coating to project a pattern. Blocked light is reflected in a different direction. Unlike a DMD, the reflective surface is static while the liquid crystals determine the reflection of light [3].LCoS devices possess many of the same advantages over LCD devices as DMD devices. LCoS devices generally have higher fill rates and smoother reflective surfaces than the DMD.

LCoS devices generally have higher fill rates and smoother reflective surfaces than the DMD. Unfortunately LCoS chips have poor contrast ratios and do not produce deep blacks. InMPµSLA, projected images will have a base level of intensity even in areas that are meant to be absent of light. This may unintentionally cure photopolymer and may even encourage the development of artifacts or otherwise restrict dimensional accuracy within the cross section. Furthermore, LCoS chips are difficult and expensive to manufacture making their availability and cost to performance ratio less than that of DMDs [27].3.4 Imaging OpticsThe optical components that modify the projected image are regarded as the imaging optics. This series of lenses is typically designed with the intention of capturing the light projection transmitting from the pattern generator and focusing it on the build surface.

This series of lenses is typically designed with the intention of capturing the light projection transmitting from the pattern generator and focusing it on the build surface. By choosing lenses with specific focal lengths and numerical apertures, a desirable reduction ratio can be achieved.The reduction ratio is what determines the achievable resolution and working area, but there is a tradeoff between the two characteristics. A higher reduction ratio (lower magnification) means that higher resolutions are achievable, but this reduces the overall projection area and results in smaller part sizes. System magnification can be calculated using lens equations. For example, Choi explains for a system with a magnification of 0.434, one pixel on the DMD (pitch of 13.68 µm) would reduce to approximately 5.9 µm on the resin surface [17].

For example, Choi explains for a system with a magnification of 0.434, one pixel on the DMD (pitch of 13.68 µm) would reduce to approximately 5.9 µm on the resin surface [17]. This constitutes the physical limit of the X-Y resolution of a system, provided that the recoat method or print orientation does not inhibit achievable feature sizes.3.5 RecoatThere are several methods for reapplying resin to the build surface. The methods includerecoating by gravity, by spreading, by pumping and by dipping (Figure 3).in microstereolithography. In this method, the build platform is dipped below the resin surface to exactly one layer thickness depth. That layer is cured, and the stage descends into the vat of photopolymer, which allows for fresh uncured resin to flow over the previously created layer.

That layer is cured, and the stage descends into the vat of photopolymer, which allows for fresh uncured resin to flow over the previously created layer. The stage then returns to a location where the previously cured layer is exactly one layer thickness below the resin surface [4–8], [17–20].Recoat by dippingisthe most commonly used methodforrecoatingRecoat by gravity is used exclusively in bottom-up projection systems (explained further in section 3.6). Unlike the dipping process when projecting from above, the build platform is raised by the z-axis actuator to create a gap between the platform and the resin container for the next layer. If the photopolymer has a low viscosity, it will flow into the created gap. If the photopolymer has a high viscosity, the platform needs to be raised to an exaggerated height such that the photopolymer can fill the gap.

If the photopolymer has a high viscosity, the platform needs to be raised to an exaggerated height such that the photopolymer can fill the gap. The z-axis actuator than moves the platform down until the desired layer thickness gap is achieved, while extra photopolymer is forced to the sides [26]. Hypothetically, pumping the resin into the created gap could assist the recoat process when the viscosity is too high.Recoat by pumping is used rarely and does not refer specifically to a recoat practice. Systems that utilize pumping in the recoat process typically are paired with a dipping process. One instance of pumping identified in literature is a syringe pump that purges uncured resin from the build surface with a high density “filler” material to prepare for the next layer [21].

One instance of pumping identified in literature is a syringe pump that purges uncured resin from the build surface with a high density “filler” material to prepare for the next layer [21]. In a multi-material system developed by Choi, a “deep dip” process is imitated through precise and extensive pumping [25].Wiping or spreading uncured resin for recoating is a method common in stereolithography practices [28], but is seldom used in microstereolithography. However, some MPµSL systems utilize a squeegee or wiping system to recoat or level the resin surface [22].3.6 Projection OrientationWhen deciding upon a base design for a microstereolithography system, besides the dynamic mask, one of the most critical considerations is the orientation in which the system will be projecting light on the build surface.

Typically, there are two orientation options as shown in Figure 4: a top-down approach where the path of light is projected from above onto the resin surface (also referred to as “from above” projection orientation) and a bottom-up approach where light is projected through the bottom surface of vat through a transparent window (also referred to as “from below” projection orientation).Figure 4. Illustration of different projection orientations: (a) image projection from above and (b) image projection from below.3.6.1 Projecting from aboveTypically the slower and more common of the two possible system orientations, projecting from above generally utilizes the dipping method to recoat the part. This approach demands that a designer account for resin characteristics, such as surface tension, viscosity, and wetting, to tune the process and achieve the desired layer thickness.

This approach demands that a designer account for resin characteristics, such as surface tension, viscosity, and wetting, to tune the process and achieve the desired layer thickness. In addition, these characteristics affect polymer settling time, which along with actuator speed, directly influences the process throughput.In an effort to reduce recoat times, vibration assisted leveling has been experimented with to encourage quicker leveling time of the photopolymer and to obtain thin layers [20], [29]. Also, as mentioned in Section 3.5, Takahashi provides one instance in which a squeegee mechanism is used for adjustment of the resin surface in MPµSL [22].

Also, as mentioned in Section 3.5, Takahashi provides one instance in which a squeegee mechanism is used for adjustment of the resin surface in MPµSL [22]. However, as high resolution is generally in the scope of the top-down design, this type of recoat is avoided as sweeping motions may be enough to agitate the resin or catch the part, destroying the print entirely.3.6.2 Projecting from beneathProjecting images through the bottom of the photopolymer container offers severaladvantages over projecting images from above the container.Less photopolymer is required in the reservoir. Because the build part is not submerged in photopolymer as when projecting from above, the reservoir container is independent of part height and can be shallower. Less photopolymer can save time and/or reduce cost. • System complexity decreases while the overall vertical print speed increases. When projecting from below, gravity is used to more quickly move and settle the photopolymer.

• System complexity decreases while the overall vertical print speed increases. When projecting from below, gravity is used to more quickly move and settle the photopolymer. Therefore it is unnecessary to implement recoating mechanisms like squeegees or long waiting periods for photopolymer to settle.Thinner layers are theoretically possible. Achievable layer thickness in this orientation is determined by the gap between the previously printed layer and the floor of the container, which is directly limited and controlled by the resolution of the vertical actuator of the build platform.Photopolymer can polymerize faster because it is removed from ambient oxygen that would otherwise inhibit crosslinking.As photopolymer cures between the floor of the resin container and the previous layer, it adheres strongly to both. Forces upwards of 50N must be applied to overcome this adhesion and move the build platform, destroying smaller features in the process [26].

Forces upwards of 50N must be applied to overcome this adhesion and move the build platform, destroying smaller features in the process [26]. Therefore, there is an inherent tradeoff between printing speed and printing resolution for MPµSLA machines.Researchers have developed a peeling process that is used when projecting underneath to alleviate feature destruction caused by separation. In this process either the build tray or resin vat is separated from the other by applying a gradual peeling force (like tape from a table). This process reduces the separation force greatly, preserving features smaller than using the same system without peeling. Unfortunately, the peeling process may prolong total build time and negate the inherent speed advantage. Chen’s research group, however, has developed a fast mask projection stereolithography process that bypasses this design tradeoff by applying a flexible silicone membrane to the floor of the resin container [26].

This SYLGARD silicone gel maintains a thin oxidation layer directly on its surface, inhibiting photopolymerization so that a substantially smaller adhesion force is developed. The membrane is also flexible so that the polymer container can slide easily under the build part, to another fresh section of photopolymer (in the same container) for the next layer to cure. This process emulates a peeling step but with much smaller displacements and applied forces, making the impact on overall print speed negligible.4. Categorization and Analysis of MPµSL SystemsThe morphological matrix introduced in Figure 3 provides a basis on which existing MPµSL systems can be analyzed. The existing systems are presented in this section via a categorization based on the critical architectural components of each system. Each morphological matrix is accompanied by a corresponding table that specifies the systems’ respective performance parameters.

Each morphological matrix is accompanied by a corresponding table that specifies the systems’ respective performance parameters. Together, this data provides a basis for determining the effect of system design decisions on critical performance parameters.4.1 First Generation – LCD MPµSL SystemsBertsch, Chatwin, and Monneret published the first MPµSL systems. These systems all use beam expansion for their conditioning optics, LCD masks to pattern light, project images from above the photopolymer, reduce the projected image in size, and recoat photopolymer by dipping. This design architecture is illustrated in Figure 5 using the morphological matrix previously present in Section 2.2, performance parameters are included in Table 2.Figure 5.

This design architecture is illustrated in Figure 5 using the morphological matrix previously present in Section 2.2, performance parameters are included in Table 2.Figure 5. Morphological matrix categorizing MPµSL machines that use a LCD pattern generator to project images from above the build platform.Table 2: Performance parameters for the MPµSL systems categorized in Figure 5.Ref[1], [2][10–13][14]Research GroupBertschChatwinMonneretYear199719981999Special Features LCD,Laser, 515 nmLCD,Laser, 351.1 nmLCD,Lamp, 530nmMin. Feature Size5 µm5 µm5 µmMin. Layer Thickness5 µmNot reported10 µmMaximum Part Size 1.3 x 1.3 x 10 mm350 x 50 x 50 mm33.2 x 2.4 x 1.3 mm3Vertical Build Rate 110 layers in 90 minutes 60 second exposure per layer 60 layers per hourAll of these machines are capable of producing 5 µm minimum feature sizes with layer thicknesses ranging from 5 to 10µm.

The machines published by Bertsch and Monneret are particularly noteworthy because they use visible light to cure photopolymer. Future MPµSL research moved away from visible photopolymers in favor of curing with UV light. Exposure times for these machines varied, but were roughly 60 seconds per layer on average. The following section introduces second generation MPµSL systems.4.2 Second Generation – DMD MPµSL SystemsThe systems categorized in this section use a DMD to dynamically shape the projected image. These systems were grouped together in order to identify the paradigm shift in design decisions with the advent of the DMD. As outlined in Section 3.3.2, the transition from LCD to DMD was a result of better performance characteristics such as smaller pixel sizes, narrower gaps between pixels, higher modulation speeds, and high reflectivity.

While smaller feature sizes were achieved with DMD based systems, the average achievable feature size is not significantly smaller than LCD systems. Unfortunately, the vertical build rate metric for many of these systems are not comparable as many publications use different speed metrics (as discussed in Section 2.3). Regardless, all systems in Figure 5 and Figure 6 use the dipping recoat method, and their build times seem comparable.

Regardless, all systems in Figure 5 and Figure 6 use the dipping recoat method, and their build times seem comparable. However, as LCD masks transmit only 12.5% of UV light (as mentioned in Section 3.3.1), build time can increase by a factor of 7 when compared to systems using a DMD with the same light source.Figure 7 and Table 4 presents systems using a DMD dynamic mask to project images fromabove, while employing alternative solutions to achieve recoating.Figure 7: Morphological matrix categorizing MPµSL machines that use alternative recoat methods.Table 4: Performance parameters for the MPµSL systems categorized in Figure 7.Ref[25]Research GroupWickerYear2009Special Feature DMD, 365 nm, multi materialMin. Feature Size~50 µmMin.

Feature Size~50 µmMin. Layer Thickness21 µmPart Size ~ 2 x 2 x 4 mm3Vertical Build Rate8-12s per layer[22]Takahashi2000DMD, 365 nm50 µm200 µm~ 2 x 2 x 2 mm3not reported[21]Roy2008DMD, 355 nm50 µm50 µmnot reported60s per layerComparing the systems’ performances listed in Tables 2 and 3 versus Table 4 (which primarily differ by recoat method) suggests that systems that do not use recoat by dipping generally have larger minimum feature sizes (approximately 50 µm) and layer thicknesses ranging from 21 µm to 200 µm. This indicates that dipping may be the preferable method forrecoating if the goal of system is to achieve a minimum feature sizes (as discussed in Section 3.5).4.3 Third Generation – From Below ProjectionAs the embodiment of second generation systems became a standard, some researchers have begun experimenting with alternative projection orientation in the effort to reduce build time.

The system architecture presented in Figure 8 and Table 5 is unique as it projects from below and recoats by gravity.Figure 8. Morphological matrix categorizing MPµSL machines that use a DMD pattern generator to project images from below the build platform.Table 5. Performance parameters for the MPµSL systems categorized in Figure 8.Ref[26][24]Research GroupChenKangYear20122012Special Features DMD, Lamp, visible DMD, Lamp, UVMin. Feature Size400 µmNot reportedMin. Layer Thickness100 µmNot reportedMaximum Part Size 48 x 36 mm2 14.6 x 10.9 mm2Vertical Build Rate 180 mm per hourNot reportedThe feature sizes reported in Table 5 are much larger than those of the systems listed inthe Tables 2-4, but the part size is also much larger. The print speed is the quickest of all systems analyzed [26]. This suggests that projecting from below is indeed quicker, but cannot achieve the resolution of top down projection systems.

The print speed is the quickest of all systems analyzed [26]. This suggests that projecting from below is indeed quicker, but cannot achieve the resolution of top down projection systems. Work in this system embodiment still developing, and there is room for further improvement based on the performance parameters advantages.4.4 Alternative EmbodimentsThe last system is unique in that it is the only published system that utilizes an LED light source, and is the only published system that uses a LCoS device to digitally pattern light. Figure 9 and Table 6 outline this system and characterizes its performance.Figure 9. Morphological matrix categorizing MPµSL machines that use a LCoS pattern generator to project images from above the build platform.Table 6. Performance parameters for the MPµSL systems categorized in Figure 9.Ref[3]Research GroupZhengYear2012Special Features LCoS, LED, 395nmMin. Feature Size1.3 µmMin.

Performance parameters for the MPµSL systems categorized in Figure 9.Ref[3]Research GroupZhengYear2012Special Features LCoS, LED, 395nmMin. Feature Size1.3 µmMin. Layer Thickness10 µmMaximum Part Size 2.56 x 1.44 mm2Vertical Build RateNot reportedThis unique system achieves very small feature sizes, comparable layer thickness and part size of other projection from above, recoat by dipping systems. It follows the trend that high resolution is achievable by combining recoating via dipping with the top-down orientation.5. ClosureThis paper presents a MPµSL morphological matrix that is designed to functionally categorize and compare the fundamental design decisions solved in the realization of MPµSL systems. Using this matrix, several published MPµSL systems are analyzed in terms of critical MPµSL performance parameters: minimum feature size (µm), layer thickness (µm), build volume (mm3), and vertical build time (mm/s).

These performance parameters are expressed in units that serve as a standard benchmark to promote direct quantitative comparisons between MPµSL systems. From this analysis, relationships between system performance and the corresponding subsystems solutions have been identified to indicate general trends and tradeoffs.The most common system embodiments project images from above the photopolymer container and dip the build platform to achieve recoating. These systems are capable of achieving the smallest feature sizes, which are close to 1 µm.Dynamic mask selection does not significantly affect the minimum achievable feature size. While pixel pitch is a relevant design consideration, final feature size is ultimately determined by the imaging optics.Dynamic mask selection significantly affects the vertical build rate of a MPµSL system.

LCD masks transmit only 12.5% of UV light, which can increase build time by a factor of 7 when compared to using a DMD with the same light source.The relationship between part size and feature size is equivalent to the relationship between aspect ratio and pixel pitch for any dynamic mask. Future dynamic masks with an equivalent pixel pitch but with greater resolution will enable creation of smaller feature sizes on equally sized parts, or the same feature sizes on larger parts. If a fast vertical build rate is priority, then system embodiment should project images from below the photopolymer vat. By using a gravity-assisted recoat approach, vertical build rates can be improved by an order of magnitude.The type of light source is noncritical provided that emission properties (wavelength and intensity) are suitable, and appropriate light conditioning is performed.Amongst published MPµSL systems, design trends have changed historically.

Exploration of novel form factors and embodiments has been limited. An area for future research is in exploring more unique embodiments and subsystem solutions. In doing so, the relationships between MPµSL performance parameters and subsystem solutions can be more broadly and quantitatively compared. Ultimately, such relationships could be used to develop processes more optimized for desired performance as governed by the application.6. AcknowledgementsEarl Campaigne is supported by a research assistantship provided by a grant from the Center for Innovative Technology. Philip Lambert is supported by a research assistantship provided by Virginia Tech’s Institute for Critical Technology and Applied Science.References[1] A. Bertsch, S. Zissi, and J. Y. Je, “Microstereophotolithography using a liquid crystaldisplay as dynamic mask-generator,” 1997.[2] A. Bertsch and J. C.

SIGN ENHANCEMENT OF BIOMEDICAL SCAFFOLDS MADE BY SELECTIVE LASER MELTINGM. Speirs*, G. Pyka†, J.-P. Kruth*, J. Luyten†, J. Schrooten†, M.Wevers† and J.Van Humbeeck†University of Leuven (KU Leuven), Department of Mechanical Engineering, Celestijnenlaan 300B, 3001 Leuven, Belgium† University of Leuven (KU Leuven), Department of Materials Engineering, Kasteelpark Arenberg 44, 3001 Leuven, BelgiumAbstractSelective laser melting (SLM) is increasingly used to fabricate biomedical scaffolds. However, the intrinsic specifications of the process such as laser spot size, layer thickness, and particle size limit the production accuracy, altering the geometrical characteristics and mechanical properties of the scaffolds. This work attempts to assess and improve the mechanical properties of TiAl6V4 biomedical scaffolds by eliminating/modifying the sharp and thin nodes (as the main source of stress concentrations and lowering the mechanical properties).

This is carried out through a gradual increase of the beam (strut) thickness around the nodes where corresponding struts meet. The compression performance of these scaffolds was assessed and compared to common examples (unaltered struts) and to scaffolds designed with thicker struts in the centre of the beams (demonstrating the largest contrast). The findings prove that the thickening of the nodal points improves the strain distribution while maintains the mechanical properties at an identical solid volume fraction. This can be used to improve the scaffold design by a gradual strut thickness (in a comparable volume fraction) for an improved bio-mechanical performance.1. IntroductionSelective laser melting (SLM) is an additive manufacturing (AM) technique which fabricates three dimensional parts layer by layer from a defined CAD model. Thin layers of powder are melted upon one another locally with a laser beam. The process is repeated for subsequent layers until completion [1].

Thin layers of powder are melted upon one another locally with a laser beam. The process is repeated for subsequent layers until completion [1]. The layer-wise nature of the process allows production of complex customised parts such as biomedical scaffolds. An overview of the process is shown in figure 1.Laser sourceScanner mirrorsPolymer coaterf-θ lensMelt poolInert atmosphereCAD ModelFeed containerBuild containerManufactured part1042Figure 1. Schematic illustration of the selective laser melting process (SLM).Scaffolds are porous materials which assist in the healing of large bone defects. Theuse of scaffolds is to allow cell seeding and eventual vascularization of the implant while providing load bearing support. The mechanical properties are dependent on the material and on the geometrical features of the scaffold [2,3].

The mechanical properties are dependent on the material and on the geometrical features of the scaffold [2,3]. For bioresorbable and organic scaffolds the main focus is maximum surface area allowing minimal implant integration time with the implants volume fraction a minor consideration. For permanent metallic scaffolds there is no option for removal. Therefore, generating an implant with sufficient open porosity and minimal solid volume fractions and adequate mechanical properties is the key factor. Ti6Al4V is a common bio-compatible material with wide usage (enabling the comparison of the findings with other work).So far, most work has focussed solely on production of uniform scaffolds with constant strut thickness and analysing methods of those products [4]. However, no attempt has been made to implement any gradual changes into the scaffold design in order to reinforce nodal zones against rupture/failure for load bearing applications.

However, no attempt has been made to implement any gradual changes into the scaffold design in order to reinforce nodal zones against rupture/failure for load bearing applications. This study aims to assess the failure modes seen in scaffold production and to improve the mechanical properties (at a specific solid volume fraction) through reinforcing the selected zones of the scaffolds. A diamond unit cell is used with two variants. One is designed with enlarged nodal joints, in contrast to the one with increased thickness within the centre of the strut. They were designed at comparable overall solid volume fractions with three different strut sizes chosen for each design. These designs were also selected to allow comparison with previous work [2].2. Materials and methods2.1 Design and productionSelective laser melting was used for manufacturing porous titanium (Ti6Al4V-ELI) scaffolds using a diamond unit cell with two variations.

Materials and methods2.1 Design and productionSelective laser melting was used for manufacturing porous titanium (Ti6Al4V-ELI) scaffolds using a diamond unit cell with two variations. One variation introduced a gradual increase in strut thickness from the centre of the beam to the node at a ratio of 2:1. The other variation incorporated the opposite with a gradually increasing thickness from the node to the centre of the beam at a ratio of 2:1. These two designs will be referred to as centrally thickened beam (CTB) and reinforced nodes (RN) along with the conventional diamond unit cell with constant strut thickness (Di): figure 2. Each design was created using Magics software [Materialise NV, Haasrode, Belgium].RN CTB DiA)B)C)D)1043Figure 2.

Each design was created using Magics software [Materialise NV, Haasrode, Belgium].RN CTB DiA)B)C)D)1043Figure 2. CAD images of each beam for each design and the pore and strut size for one unit cell: A) different beam types (left to right as RN, CTB, Diamond), B) reinforced node (RN), C) centrally thickened beam (CTB) design, and D) conventional diamondAll scaffold samples were designed with a height of 6.4mm and 6mm diameter with aconstant pore size of 1000µm and varying strut sizes as shown in table 1. All samples were removed from the SLM base plate by EDM and then ground to a height of 6mm.Table 1.

All samples were removed from the SLM base plate by EDM and then ground to a height of 6mm.Table 1. Overview of scaffold designs producedScaffold DesignStrut diameter (µm)Surface area (mm2) Volume (mm3) Designed VolumeDiamond14036313.0fraction (%) 7.716040116.49.718043019.711.620045923.413.8Centrally Thickened Beam100-200120-24033837712.216.47.29.7140-28044322.313.2Reinforced Nodes100-20031010.96.4120-24035215.08.814028042321.412.6The designed scaffolds were produced on an in-house developed SLM machine using an IPG Yb:YAG fibre laser of 300W with Ti6Al4V powder. All designs were produced with a 45µm beam offset Table 2 gives an overview of the powder specifications and SLM process parameters.Table 2.

All designs were produced with a 45µm beam offset Table 2 gives an overview of the powder specifications and SLM process parameters.Table 2. Specifications of the Ti6Al4V powder, produced bulk material, and SLM process parametersSpecifications of the Ti6Al4V powderSLM process parametersDensity: 4.43g/cm3 Laser power: 42W Average grain size: 10-45μm E-modulus: 110GPa Tensile Strength (bulk material): 1000MPa Yield stress at 0.2% deformation: 920MPaScanning speed: 260mm/s Hatch spacing: 74μm Layer thickness: 30μm Beam offset 45μm Beam diameter 52µm2.2 Morphological CharacterizationMicrofocus X-ray computed tomography (μCT)-based morphological characterization of the Ti6Al4V scaffolds was performed using a Phoenix NanoTom S (GE Measurement and Control Solutions, Wunstorf, Germany) with a 180kV / 15 W high performance nanofocus X- ray tube and a 2304x2304 pixel Hamamatsu detector. A voltage of 90 kV and 240 μA were applied.

A voltage of 90 kV and 240 μA were applied. A tungsten target was used and a 0.3mm cupper filter was installed. Each sample was scanned over 360º with a 0.15° rotation step, an exposure time of 500ms with no image skip1044and no frame averaging was applied. After scanning reconstruction was completed using Phoenix datos/x 2.0 reconstruction software (GE Measurement and Control Solutions, Wunstorf, Germany). The resulting images had an isotropic voxel size of 6.5 μm. CTAn software (Bruker micro-CT, Kontich, Belgium) was applied for 3D morphological analysis of the µCT data. Automatic Otsu segmentation [5] was applied for binarization of the reconstructed µCT images. For each geometry 3 samples were scanned. Additionally, an experimental strain analysis of one CTB and one RN sample was performed based on the non- rigid image registration of the µCT images [6-7]Each type of scaffold was viewed using scanning electron microscopy (SEM; Phillips,XL 30, Germany) operating at 10kV.

An ultrasonic cleaning in ethanol solution for 10 minutes was completed before analysis. The volume fraction and beam thickness were calculated for each design with the structure thickness calculated in 3D using a sphere fitting algorithm [8].2.3 Mechanical characterizationFor mechanical testing, three as-produced samples of each design were compressedusing an in house developed in-situ loading stage with maximum available load 30kN. A pre loading step of 0.01kN and a compression rate of 0.2mm/min was applied. The structural stiffness and ultimate compressive stress were calculated from the stress strain curves obtained from the load displacement data.For strain mapping a radio-translucent micro-mechanical compression setup is used toapply and maintain strain for high resolution μCT scanning. To evaluate the local strain changes in function of displacement a constant compression rate of 0.2mm/min was applied followed by μCT scanning.

To evaluate the local strain changes in function of displacement a constant compression rate of 0.2mm/min was applied followed by μCT scanning. A pre-load of 0.01kN was applied and then the reference scan of the non-compressed sample was applied. The sample was then compressed to 50% ultimate compressive strain (50% UCS) followed by μCT scanning. Both μCT images taken for each sample were registered to each other non-rigidly using Elastix software [8]. One CTB-140 and RN-140 sample were used for strain mapping due to their similar volume fractions post production.3. Results and Discussion3.1 Morphological and Mechanical CharacterizationDespite the SLM potential to produce biomedical scaffolds, the geometry of thescaffolds is still restricted by the SLM characteristics such as laser spot size and layer thickness in table 2. Some of these issues are shown in figure 3.

Some of these issues are shown in figure 3. For example, struts of 100µm (designed strut thickness at the centre of the strut in figure 3B) are difficult to be achieved when a laser spot size of 80µm is used while the variation of beam diameter between layers is 10μm. On the other hand, even a thin layer thickness of 30µm has led to a pronounced staircase effect on the thin struts. The residual particles attached on the scaffolds (originating from the laser energy in margins of the spot diameter) are also evident. This is even more critical at nodes where the struts meet.

The residual particles attached on the scaffolds (originating from the laser energy in margins of the spot diameter) are also evident. This is even more critical at nodes where the struts meet. The residual particles potential release post1045implantation are an issue for the application (surface cleaning by chemical/electrochemical methods are usually followed to resolve this issue) [9].In terms of design, the gradual increase of the strut was less pronounced compared to the CAD design file, once again due to working close to the machine limits, i.e., a laser spot size of 80µm used to melt thin struts (figure 3). This is even more evident for the RN scaffolds where the struts lowest diameter should be located in the centre of each strut (figure 3B).A)B)Figure 3. SEM images of produced scaffolds: A) Centrally thickened beam (CTB) and B) Reinforced nodes (RN)3.2 Mechanical properties and Micro-CT analysisThe morphological parameters were assessed by micro-CT.

The solid volume fractionwas assessed and compared to the structural stiffness as shown in figure 4 for all produced designs. The average value for the 3 samples scanned and compressed is given. From the mechanical properties it can be seen that CTB showed a much lower stiffness at similar solid volume fractions when compared to the other two designs. It appears that RN performed equally with diamond.1046Figure 4.Comparison of solid volume fraction in relation to stiffness for each designThese issues result in analytical methods for unit cells producing theoretical stiffness much higher than those observed when using constant strut thickness in other studies [10]. A relationship between the Young’s modulus and volume fraction scales on the power law and can be defined as:()Where E*= Youngs modulus of the scaffold, Es= Youngs modulus of bulk materials, P*= scaffold density, Ps=bulk density, C1=constant dependent on cell geometry [11].

Where C1≤1 and from analytical methods values of 0.98 are observed and from SLM produced Ti6Al4V C1 values of 0.1-0.3 are found in literature [2,4].3.3 Local Strain MappingFigure 5 depicts a histogram of the volumetric strain calculated experimentally forCTB-140 and RN-140. As seen, the strain analysis revealed a large variation of strain frequency dependent on the design. The strain histograms of the RN-140 structure had a different bell-shape and asymmetry with respect to the mean value in comparison to CTB- 140. In general, strain histogram analysis indicated a more homogenous strain distribution observed for RN-140 with high frequency of low compressive strains while CTB-140 revealed a wide range of compressive and tensile strains.Analysis of the mean strain indicated differences in the level of the local deformations of the CTB-140 in comparison with RN-140 (4.3% and 5.9% of compressive strain for RN-140 and1047CTB-140 respectively).

Measurements of the most frequent (dominant) strain revealed a similar pattern (7.2% and 3.0% of compressive strain for RN-140 and CTB-140 respectively).Figure 5. Frequency of local distribution strain for CTB and RN at 50% of ultimate compressive strain (50% UCS)Figure 6 represents the 2D visualization of the strain computed for CTB-140 and RN- 140 in the coronal sagittal and transverse slices. It can be shown that a more uniform strain is observed for RN-140. (Figure 6B). This confirms visually the differences in the strain histograms (Figure 5). CTB -140 shows large sections with high compressive strain (indicated in pink colour in Figure 6a). For RN-140 the largest compressive strains are observed at the centre of the beams. This is in contrast to CTB-140 where high compressive deformations were observed mainly in the connection between beams and nodes.

This is in contrast to CTB-140 where high compressive deformations were observed mainly in the connection between beams and nodes. Perhaps, thicker size of struts at central zones have led to a localized strain by further thickening of those areas in comparison with nodal joints where the struts have bended over and produced a tensile strain. In contrast, the gradual increase in strut thickness from the centre of the beam eliminated the amount of the critical points of the scaffolds. . Additionally, RN-140 revealed a more homogenous strain distribution across the scaffold structure in comparison to CTB-140 (Figure 6) which could be advantageous during implant loading triggering osteoinduction uniformly throughout the scaffold.1048A)B)Figure 6.

Typical strain maps (coronal, sagittal and transverse) of the obtained 3D μCT images of the compressed Ti6Al4V scaffold designs showing the volumetric strain A) Centrally thickened beams (CTB) and B) Reinforced nodes (RN)In figure 7 a 3D visualization of the strain maps for CTB-140 and RN-140 are given, confirming the results present in two dimensions in figure 6. Once again, it can be observed that a large inhomogeneous deformation distribution of the CTB occurs in contrast with the very homogenous distribution of the deformation within the RN design.A)B)Figure 7. A 3D visualization of the local strain at 50% ultimate compressive strength for A) centrally thickened beam (CTB) and B) reinforced nodes (RN).1049As mentioned, the gradual change of strut thickness changed the scaffold architecture (figure 3), though it was restricted by the machine limitations (e.g. high comparative size of the laser beam diameter to the strut).

high comparative size of the laser beam diameter to the strut). The RN scaffolds showed a similar stiffness to the conventional diamond design, which was higher than CTB scaffolds. A more homogeneous deformation distribution was observed for the scaffolds designed thicker at the nodes leading to failure occurring at both nodes and struts. Higher deviation in the gradual designs should be followed in future works to better clarify the alternations in mechanical performance at similar volume fractions. Applying surface modification techniques (such as chemical/electrochemical etching) to remove the attached particles on these enhanced scaffolds is another promising aspect, pursued for future work.4. ConclusionThis study assessed the strain distribution and structural stiffness for two variations of a standard diamond unit cell.

ConclusionThis study assessed the strain distribution and structural stiffness for two variations of a standard diamond unit cell. For centrally thickened beam scaffolds (CTB) a lower stiffness was observed at similar solid volume fractions along with a largely inhomogeneous strain distribution. This was due to a larger amount of critical points throughout the scaffolds structure with elevated local deformation observed by local strain mapping experiments located mainly at the beam-node connections. For scaffolds with reinforced nodes (RN), the apparent compressive stiffness was equal to that of the conventional diamond at a similar volume fraction. Moreover, the reinforced nodes led to a very homogenous distribution of strains, being a promising aspect of the design.For future work RN scaffolds could be used in conjunction with surface modificationtechniques used to remove residual grains after production [12].

This will allow compensation for material removal at the nodes resulting in more inhomogeneous strain distributions similar to those observed in CTB scaffolds.AcknowledgementsThe authors acknowledge support within the EU 7th framework programme FP7/2007-13 under Marie-Curie project grant No. 264635 (BioTiNet-ITN) and the Hercules foundation for founding the project AKUL/09/001 “Micro- and nano-CT for the hierarchical analysis of materials” and the IAP project P7/21 INTEMATE, “Multiscale mechanics of INTErface dominated MATErials”, for financial support.References[1] J.-P. Kruth, G. Levy, F. Klocke, T.H.C. Childs, “Consolidation phenomena in laser and powder-bed based layered manufacturing,” CIRP Annals - Manufacturing Technology, Volume 56, Issue 2, Pages 730-759, 2007.[2] S. Van Bael, Y.C. Chai, S. Truscello, M. Moesen, G. Kerckhofs, H. Van Oosterwyck, J.- P. Kruth, J. Schrooten, 2012.

SIGN FOR ADDITIVE MANUFACTURING CURRICULUM: A PROBLEM- AND PROJECT-BASED APPROACHChristopher B. Williams Department of Mechanical Engineering Department of Engineering Education Virginia TechCarolyn Conner Seepersad Department of Mechanical EngineeringUniversity of Texas at AustinABSTRACTAdditive manufacturing education is of key importance because unfamiliarity with AM technologies is one of the barriers to its widespread adoption. In this paper, the authors describe their efforts in Additive Manufacturing. Their courses, offered at the University of Texas at Austin and Virginia Tech, cover the science of AM as well as principles of “design for additive manufacturing.” The courses use both problem-based and project-based pedagogies to present students with opportunities to gain hands-on experience with the technologies of AM.

Examples of project activities are presented along with student feedback.to addressthis need via an undergraduate/graduate courseKeywords: Additive Manufacturing, Engineering Education, Design Education1. ADDITIVE MANUFACTURING EDUCATION1.1. MotivationThere has been significant investment in advancing the US’s manufacturing presence in the global innovation economy in the past few years. This need has been a key focus of the United States’ federal government: the “Manufacturing Enhancement Act of 2010” was created to promote investment in the manufacturing sector [1] and the Advanced Manufacturing Partnership was launched last year [2]. The President recently proposed a “National Network for Manufacturing to boost manufacturing research and workforce development [3].

The President recently proposed a “National Network for Manufacturing to boost manufacturing research and workforce development [3]. The pilot NNMI institute is focused on Additive Manufacturing (AM), further illustrating its importance to our nation’s global competitiveness.Innovation” (NNMI)There is also considerable public interest in improving US manufacturing. In a recent study by Deloitte and the Manufacturing Institute, 77% of respondents believed that the U.S. needs a more strategic approach to develop its manufacturing base, and 74% said that the U.S. should further invest in manufacturing industries [4]. However, in the same survey, it was found that the youngest respondents were the least likely to think that manufacturing is important to our economic prosperity (71%) or thought of manufacturing as “high tech” (59%).To address this fundamental disconnect, advancing manufacturing workforce development is of great importance.

The need to educate future manufacturing engineers is a core focus of the “2009 Roadmap for Additive Manufacturing,” as unfamiliarity with AM technologies is seen as a barrier to adoption. The roadmap urges the development of university courses and “programs for educating the general population to enhance the interest in AM applications and generate some societal ‘pull’ for the technologies” [5]. While some AM courses at the undergraduate and81graduate levels do exist (e.g., [6]), their limited quantity does not match the recent interest in, and suggested national importance of, the technology. To answer the call posed by the AM Roadmap the authors have designed and implemented an undergraduate/graduate AM course at their respective institutions.1.2. Design for Additive ManufacturingAM’s layer-by-layer fabrication approach removes many of the geometric constraints imposed by traditional manufacturing processes.

Design for Additive ManufacturingAM’s layer-by-layer fabrication approach removes many of the geometric constraints imposed by traditional manufacturing processes. Its capability to selectively place material allows for the fabrication of complex geometries that are specifically designed to support and improve multiple design objectives, and thus has the potential to significantly alter how products are designed. To realize the full potential of AM, engineers must know how to design products for fabrication via AM. In addition, engineers must not only understand AM technologies and materials, they must also be able to synthesize its economic and environmental impacts on a manufacturing value chain. As such, the primary objective of the authors’ courses is in advancing engineers’ ability to design and fabricate complex systems via AM.This design focus dictates the design of the authors’ courses. Design is considered to be the central or distinguishing activity of engineering [7].

Design is considered to be the central or distinguishing activity of engineering [7]. Research on engineering design thinking and learning has established that design is hard to learn and harder to teach [8]. Design instructors cannot simply “teach” design principles through lecture since the unique context of each student’s design prevents the canned delivery of common content. Such deductive modes of instruction do not reflect what engineering education research has discovered about learning and effective pedagogical practice.Instead, engineering design pedagogy follows the constructivist learning theory, in which it is postulated that students form knowledge representations of new information by building on their previous knowledge and experiences [9]. If the new information has few connections to what they already know, learning will not occur, nor will students be motivated to learn [10].

If the new information has few connections to what they already know, learning will not occur, nor will students be motivated to learn [10]. Thus, effective instruction must provide experiences in which students actively construct knowledge by adjusting, rejecting, or modifying their prior beliefs and understanding based on their experiences [11]. Deductive instruction does not facilitate this mode of learning; its “skill-and- drill” approach allows students to approach learning passively, and does not challenge them to modify their prior understanding. Design educators must augment content delivery with individual coaching and mentoring of students as they progress through their design efforts [12]. They must follow an inductive teaching approach wherein students are presented problems and projects in which to apply and synthesize their prior knowledge through a design methodology.In this paper, the authors present their approach to educating future engineers about AM.

Their design-focused approach is based on problem-based and project-based pedagogies. These inductive pedagogical approaches are presented and differentiated in Section 2. A general overview of the course is presented in Section 3, and examples of posed problems and projects are presented in Section 4. Closure is given in Section 5.822. PROBLEM-BASED AND PROJECT-BASED LEARNINGFollowing calls from the National Academy of Engineering [13] and the American Society of Engineering Education [14] for more inductive teaching in engineering education, the authors’ AM course is centered on inductive problem-based and project-based pedagogies. In inductive instruction, students are to complex and realistic problems before being exposed to relevant theory and equations (Figure 1). Students then learn what is needed to solve the presented challenge [15].

Students then learn what is needed to solve the presented challenge [15]. Inductive teaching methods include inquiry-based learning, problem- based learning (PmBL), project-based learning (PjBL), case-based just-in-time teaching.firstintroducedFigure 1. Inductive and Deductive Instructional Approachesteaching, andInductive instructional approaches are “supported by the best research on learning currently available, compatible with the currently most widely accepted theories of learning, and promotive of the problem-solving skills and attitudes that most instructors would say they desire for their students” [11]. The problems/projects are designed to be representative of authentic problems, which have been shown to motivate students, maintain their interest, and actively engage them in learning [11].

The problems/projects are designed to be representative of authentic problems, which have been shown to motivate students, maintain their interest, and actively engage them in learning [11]. PBL learning approaches have been found to improve the development of critical thinking and problem-solving [11], and to enhance understanding of critical engineering concepts [16].2.1. Problem-Based Learning (PmBL)With origins in medical school programs, the central principle of the PmBL approach is that students’ fulfillment of learning objectives is accomplished through the solution of an open- ended problems, rather than through a deductive presentation of information [17]. The problem, which is carefully designed to be authentic and reflective of professional practice, serves as the motivation for learning the content.

The problem, which is carefully designed to be authentic and reflective of professional practice, serves as the motivation for learning the content. Students work in small groups to solve the problem by first identifying what they already know, what they need to know, and how and where to access the information that will assist them in solving the problem. The problems are used as an opportunity for students to acquire the desired knowledge while simultaneously enhancing their problem solving skills and their competency for self-directed learning [18].Simply providing students with an open-ended problem is not considered true PmBL. The role of the faculty is a crucial component to successfully implement PmBL. The instructor’s role is not to act as a “sage on the stage,” but instead to serve as a facilitator or guide [18]. The instructor must guide the learning process while also leading students through reflection and debriefing at the conclusion of the experience.

The instructor must guide the learning process while also leading students through reflection and debriefing at the conclusion of the experience. The facilit ator’s role is akin to that of a consultant, wherein they provide scaffolding throughout the process to the support a student’s learning and understanding.Research in PmBL has shown that students tend to acquire a more complete understanding of concepts and ideas than those who have received traditional instruction [19]. Additional benefits include an increase in student knowledge, skill acquisition and transfer, and improving positive83attitudes about the course and profession [20]. There is also evidence of an increase in students’ metacognitive skills (e.g., self-awareness of ability to learn, retrieve knowledge, etc.), which translates to an improved ability to transfer knowledge across different contexts [21].2.2.

Project-Based Learning (PjBL)PmBL and PjBL are often confused as they are both grounded in the pedagogical approach of providing students self-directed learning opportunities via an open-ended challenge. The faculty’s role is also similar, wherein scaffolding is provided through facilitation and tutoring. The design of the project statement in PjBL is similar to that of the problem statement in PmBL: the project must be open-ended and representative of the profession; although, the projects in PjBL tend to be larger in scope (and time) and thus closer to professional reality. Students work in groups to complete the project assignment in PjBL, thus providing intrinsic motivation to demonstrate better teamwork and motivations skills [22].A key difference between PmBL and PjBL is that the project of PjBL results in a concrete and explicit artifact (e.g., model, report, computer program, artifact, etc.) that can be shared and critiqued [22].

In this regard, PjBL is focused on the synthesis and application of knowledge, whereas PmBL is focused more on the acquisition of knowledge [23]. In effect, PmBL serves as a scaffolding necessary to complete the synthesis tasks of PjBL [24]. In addition, PjBL courses are supported by the theoretical basis of situated cognition. The situated cognition theory of learning notes that knowledge is contextual and is a product of the activity and situations in which it is created [25]. By providing an authentic activity in PjBL, the domain knowledge and concepts of the field are employed and allows students to act purposefully in the activity. PjBL also offers students the opportunity for cognitive apprenticeship wherein metacognitive skills and processes are learned through interaction with the project advisor [26].3.

PjBL also offers students the opportunity for cognitive apprenticeship wherein metacognitive skills and processes are learned through interaction with the project advisor [26].3. GENERAL CLASS STRUCTUREThe development of the course’s learning objectives were guided by the need to further the advanced manufacturing workforce (Section 1.1) and the stated importance of linking design and AM (Section 1.2). Specifically, once successfully completing the course, the students should be able to:Explain the capabilities, limitations, and basic principles of alternative AM technologies.  Evaluate and select appropriate AM technologies for specific design-manufacturing applications.Explain the fundamental causes of errors and irregularities in AM parts.  Apply AM techniques to a challenging design and manufacturing application.

 Apply AM techniques to a challenging design and manufacturing application.  Identify, explain, and prioritize some of the important research challenges in AM.These stated learning objectives highlight the need for students to be able to synthesize their knowledge of AM technologies’ strengths and limitations in order to successfully design and fabricate a product. These objectives provide a foundation for selecting course topics and their relative importance. A generalized list of the course’s topics is presented in Figure 2 along with the percent of class meetings (~30 meetings in total) devoted to each class topic.84Figure 2. Overall AM Course Structure and AllocationAt the beginning of the course, the instructors provide a high-level introduction to the AM technologies through the use of a functional classification framework [27].

Case studies and commercial applications of AM are also presented to motivate student learning and to situate their learning in the context of AM professional practice. Students are then introduced to the AM technologies at a more detailed level. Typically, a single AM technology (e.g., powder bed fusion, material extrusion, vat photopolymerization, etc.) is discussed during a class meeting. The instructor leads the students through the identification of the strengths and weaknesses of each technology while also exposing them to the processes’ fundamental science. Explorations of AM science is guided by the literature produced by the AM research community and an AM textbook [28]. Students are also guided through an exploration of “AM Systems Analysis,” which focuses on technology selection and cost estimation decision strategies.This coverage of AM science accounts for ~25% of the course; it provides students the technical foundation necessary to design novel AM products.

Occupying 25% of the course content, “Design for AM” class topics are focused on a structure similar to systematic engineering design methodologies proposed by [29] and [30]. Specifically, the following topics are discussed:Identifying Opportunities: In this module, students learn about the types of applications for which AM is best suited. Students are introduced to techniques for identifying AM product development opportunities and customer needs.AM Project Planning and Economics: Students learn project planning and cost estimation techniques. Specific to AM, they discuss the impact of the digital manufacturing paradigm.  AM Concept Generation: In this module, students learn concept generation techniques and learn to employ idea generators that are unique to AM (e.g., customization, low-volume production, assembly reduction, and complex geometry).AM Embodiment Design: Students learn how best to design the structure of their product for AM in this module.

Considerations of AM tolerancing for various part features (e.g., through holes, snap-fits, living hinges, etc.) are explored.AM Detailed Design: Content in this module is focused in AM common build strategies (and potential errors) caused by part orientation, poor interlayer bonding, and resolution limitations of various AM technologies.As discussed in Section 2, the design focus of the authors’ courses mandates an inductive approach to instruction. The authors employ various PmBL and PjBL activities throughout the85semester to engage students in each of the course topics. Examples of these activities are detailed in Section 4.4. PROBLEM AND PROJECT EXAMPLES4.1. Problem-based Learning: Design of Benchmarking PartThe first PmBL activity of the course is focused on the learning objective of enabling students to explain capabilities and limitations of AM technologies (Section 3).

Instead of rote learning AM process specifications, students actively explore the technology limitations by designing and measuring a part for benchmarking one of three metrics: resolution, accuracy, or surface finish. The part is to be designed to enable students to observe the effects of potential sources of AM build error (e.g., orientation, presence of support material, layer thickness, etc.) on the chosen metric. Example student work is presented in Figure 3.(a)(b)(c)Figure 3. Example Student-Designed Benchmarking Parts; (a) Accuracy, (b) Resolution, (c) Surface FinishThe students work in small teams (3 people) to respond to the challenge.

Example Student-Designed Benchmarking Parts; (a) Accuracy, (b) Resolution, (c) Surface FinishThe students work in small teams (3 people) to respond to the challenge. The instructor guides the students through (i) defining their assigned metric (e.g., “What’s the difference between resolution and accuracy?”), (ii) identifying appropriate metrology techniques, and (iii) designing the part to measure the effects of all potential sources of build error.Once the part is designed, each of the team’s submissions is built on a Laser Sintering machine, a PolyJet 3D Printing machine, and a Fused Filament Fabrication machine. Each team then performs the appropriate measurements and analysis on each of the three parts to identify the capabilities of each AM process.

Each team then performs the appropriate measurements and analysis on each of the three parts to identify the capabilities of each AM process. Finally, the students present their work to the class in order to educate their peers (i.e., thus the team assigned to study resolution learns about accuracy and surface finish).This PmBL example provides an opportunity for students to actively investigate sources of AM process build error. In addressing the proposed challenge, students also learn metrology techniques, and abstract AM processes’ strengths and limitations. In addition to fulfilling a core learning objective of the course, it also provides a useful foundation (and motivation) for learning AM process selection techniques (Section 4.2). Finally, students often return to their results while completing their semester AM design project (Section 4.3).864.2.

Finally, students often return to their results while completing their semester AM design project (Section 4.3).864.2. Problem-based Learning: Dissection/SelectionThis assignment provides students an opportunity to investigate the advantages and limitations of available AM technologies, compare the capabilities of those technologies to conventional manufacturing processes, and practice selecting an appropriate technology for a specific application.To begin the assignment, each student selects a mechanical consumer product of interest to him/her. The product is required to be of low to moderate complexity (20-40 parts) and disassemblable with standard tools available to the students. Each student disassembles his/her product, takes photographs of exploded views of the disassembled product, and creates an accompanying bill-of-materials complete with the name of the part, its function, dimensions, mass, and likely manufacturing process and material.

Students often perform background research to identify the most likely constituent materials and manufacturing processes, especially when the material is not labeled (e.g., with a plastic recycling number) and the manufacturing process is not evident (with injection molding and sheet metal stamping being two of the most common and easily recognizable processes).Next, each student is required to identify two components from the BOM that are fabricated with significantly different manufacturing methods. Students often choose an injection-molded plastic part and a metal part that is either cast or fabricated via sheet metal stamping. For each component, students must establish a set of criteria for evaluating the quality of the component (e.g., dimensional accuracy, strength). At the beginning of the assignment, the class and the instructor construct a list of candidate criteria, including dimensional accuracy, surface finish, tensile strength, and cost.

At the beginning of the assignment, the class and the instructor construct a list of candidate criteria, including dimensional accuracy, surface finish, tensile strength, and cost. A target value is established for each criterion for each component, based on values that are reasonable for the parent manufacturing process.Finally, each student constructs a selection decision matrix to identify the most appropriate AM technology for fabricating the part in small lot sizes. Populating the decision matrix requires knowledge of the capabilities of each candidate AM process/material combination, including the typical accuracy of each process and the expected material properties (e.g., tensile strength, density, Young's modulus, operating temperatures) associated with each process/material combination. Students compile this information in a chart as each AM process is discussed in class.

Students compile this information in a chart as each AM process is discussed in class. Each student summarizes his/her recommendations, including a discussion of the strengths and shortcomings of the selected AM processes in a written report and brief classroom show/share discussion. Students are required to evaluate the sensitivity of their conclusions to shifting priorities; e.g., would the selected AM process be different if aesthetics were emphasized over strength?This assignment requires students to traverse the entire range of Bloom's taxonomy. They accumulate data on the capabilities of leading AM process/material combinations and organize it into a table that offers side-by-side comparisons of the alternatives. They apply their knowledge of manufacturing processes and materials to identify the most likely material and fabrication route for components in their product.

They apply their knowledge of manufacturing processes and materials to identify the most likely material and fabrication route for components in their product. Many students find this task much more difficult than expected, because it sometimes requires considerable knowledge of different processes and the strengths and capabilities of each one and sophisticated reasoning about the most likely fabrication path. Selecting the most suitable AM process/material for reproducing the part (and87justifying that selection) requires students to evaluate the capabilities of each alternative very carefully. By the end of the assignment, they begin to think less like a naïve customer and more like a service bureau advisor making a recommendation on a particular part.4.3. Project-based Learning: Semester Design ProjectThis project provides students an opportunity to learn how to design for AM by creating a new product that is uniquely suitable for AM.

Project-based Learning: Semester Design ProjectThis project provides students an opportunity to learn how to design for AM by creating a new product that is uniquely suitable for AM. These actions represent the highest levels of Bloom’s Taxonomy of learning [31].In small groups of 3-5 participants, students identify, design, fabricate, and test a product that is suitable for AM. When selecting a design problem to solve, students are required to identify compelling technical and/or economic reasons for additively manufacturing the product, rather than utilizing more conventional manufacturing techniques. For example, AM is often driven by the need for customization, rapid cycle times, and/or small lot sizes; economic viability is also important to consider.Once the design opportunity is identified, students proceed to further clarify the task and develop concepts. Each group generates a formal problem statement and a requirements list (or specifications sheet).

Each group generates a formal problem statement and a requirements list (or specifications sheet). They also design and implement a procedure for customer needs analysis and concept generation, leading to a set of preliminary concepts (often in the form of hand sketches). Each group refines and selects a leading concept for embodiment and detailed design. They perform detailed technical and economic analysis of the chosen concept by, for example, conducting FEA analysis of the mechanical performance of the product and economic analysis of the cost of fabricating the product at a service bureau. They iteratively refine the design as necessary, culminating in a CAD model and .stl file of the final design. The instructional team fabricates the part with in-house AM machines. After receiving their fabricated parts, students design and implement a testing and evaluation procedure for assessing the technical performance of the fabricated product (as well as its usability, if applicable).

Finally, students critically evaluate their results and document their projects in the form of final presentations and reports.Figure 4 illustrates a selection of final projects from the UT Austin class. Many students pursue products that are customized for individual users. For example, the back support in the upper left is designed to be a form-fitting support between a chair and the back of a specific student with recurring back pain. A clay mold of the student’s back was generated and converted to an equivalent CAD surface with a 3D scanner. Similarly, the tennis racquet handle in the lower right is customized for the grip of a female student. The car radio compartment insert in the upper right is customized with an individual student’s name and personalized design, and it is tailored to hold an iPod and fit snugly into the car radio slot in a typical dashboard. Other groups choose to make functional retail products with highly complex or specialized shapes or topologies.

Other groups choose to make functional retail products with highly complex or specialized shapes or topologies. The gear shift handle (bottom row of Figure 4, second from right) assumes an organic shape that fits the palm and fingers of a typical user ergonomically. The desktop novelty clock in the middle of the top row includes a path for a marble to wind into and out of the body of the clock, along with customized geometric shapes around the outside of the clock. Other students build single-lot prototypes with special features, such as the energy harvesting device in the bottom row (second from left). The device houses inductive energy harvesting components that can be attached to a bridge. Rotating the top cap of the device tunes88the harvester to the fundamental frequency of vibration of the bridge. The device replaces an experimental prototype that required more than 40 additional parts and twice the bounding volume.

The device replaces an experimental prototype that required more than 40 additional parts and twice the bounding volume. The device on the bottom left is intended to be a small- or single-lot production device for teaching and outreach. It is comprised of rotating elements to represent many of the facets of mechanical engineering, including moving gear trains (fabricated in their assembled state), a radial engine with moving pistons, a biomechanical leg with a flexing knee, and a rotating turbine.At UT Austin, the semester culminates with a final presentation and a best-of-show award sponsored by a local company. Select projects are displayed in hallway display cases. Many students find the project both challenging and enjoyable. Unlike other design projects in the curriculum, students must adopt a very entrepreneurial perspective to identify and justify an appropriate AM-related design opportunity. The design problem is not provided or strictly bounded.

The design problem is not provided or strictly bounded. They also follow the product development process from idea inception to fabrication and performance verification. Students often learn important lessons about design for AM when their products take unexpected forms, such as features that mate imperfectly, walls that are too thin to be watertight, or gears that either fuse or mate so loosely that they frequently slip.Figure 4. A selection of final projects from the AM course at UT Austin5. CLOSURE AND FUTURE WORKIn this paper, the authors present their work towards designing an Additive Manufacturing course. Given its central role in the profession of engineering, the crux of the course is the topic of Design for Additive Manufacturing. As engineering design pedagogy follows the constructivist learning theory, the authors chose an inductive instructional approach to their course.

As engineering design pedagogy follows the constructivist learning theory, the authors chose an inductive instructional approach to their course. In this approach, students are presented challenging problems (Problem-based Learning) or open-ended projects (Project-based Learning). In addition to promoting a desired skill-set, this pedagogical approach aligns with what engineering education research has discovered about learning and effective pedagogical practice. Specifically, inductive approaches89provide students an opportunity to learn experimental skills, mirror professional practice, and build and test designs, while encouraging discovery and independent learning, and improving motivation, teamwork, and communication [32].The authors present three sample activities in this work. In one activity, students are challenged to design, measure and analyze a benchmarking part for understanding AM part surface finish, resolution and accuracy (Section 4.1).

In one activity, students are challenged to design, measure and analyze a benchmarking part for understanding AM part surface finish, resolution and accuracy (Section 4.1). Students learn how to evaluate and select Additive Manufacturing processes for product fabrication via a problem-based activity centered in product dissection. Finally, students apply their newly-acquired knowledge of design for AM by carrying a final project through the entire product development process from idea inception to AM to testing.Anecdotally, the authors have received significant positive feedback from their students in this course. The authors look next to formalizing an assessment strategy for exploring student learning as a result of the class. The authors will investigate how this course affects students’ opinions of advanced manufacturing, self-efficacy, and metacognition.6.

The authors will investigate how this course affects students’ opinions of advanced manufacturing, self-efficacy, and metacognition.6. ACKNOWLEDGEMENTSThe authors would like to acknowledge the support provided by Harvest Technologies and the Laboratory for Freeform Fabrication at UT Austin, including the assistance of Mark Phillips who fabricated the parts. The Virginia Tech Department of Mechanical Engineering is acknowledged for supporting laboratory costs for the Virginia Tech course. The authors acknowledge Dr. David Rosen (Georgia Tech) for providing the inspiration for the courses discussed in this paper. The authors would also like to thank the students of the described courses for their enthusiastic participation.7. REFERENCES1. Associated Press, 2010, August 11, "Obama signs manufacturing bill," The Washington Times.2.

REFERENCES1. Associated Press, 2010, August 11, "Obama signs manufacturing bill," The Washington Times.2. White House Press Secretary, 2011, "President Obama Launches Advanced Manufacturing Partnership," http://www.whitehouse.gov/the-press-office/2011/06/24/president-obama- launches-advanced-manufacturing-partnership.3. The White House, Office of the Press Secretary, March 9, 2012, "President Obama to Announce New Efforts to Support Manufacturing Innovation, Encourage Insourcing," http://www.whitehouse.gov/the-press-office/2012/03/09/president-obama-announce-new- efforts-support-manufacturing-innovation-en. Development LLC, 2009, "Public Viewpoint on Manufacturing," http://www.deloitte.com/dtt/cda/doc/content/us_mfg_manufacturingviewpoint060809.pdf .5. Bourell, D., M. C. Leu and D. W. Rosen, Eds. (2009). Roadmap for Additive Manufacturing: Identifying the Future of Freeform Processing, The University of Texas at Austin. 6. Bohn, J.

Bourell, D., M. C. Leu and D. W. Rosen, Eds. (2009). Roadmap for Additive Manufacturing: Identifying the Future of Freeform Processing, The University of Texas at Austin. 6. Bohn, J. H., 1997, "Integrating Rapid Prototyping into the Engineering Curriculum - A Case Study," Rapid Prototyping Journal, 3, 1, 32-37.907. Simon, H. A., 1996, The Sciences of the Artificial, MIT Press, Cambridge, MA. 8. Atman, C., D. Kilgore and A. McKenna, 2008, "Characterizing Design Learning: A Mixed- Methods Study of Engineering Designers’ Use of Language," Journal of Engineering Education, July, pp. 309-326.9. Pellegrino, J. W., 2006, "Rethinking and redesigning curriculum, instruction and assessment: What contemporary research and theory suggests," National Center on Education and the Economy for the New Commission on the Skills of the American Workforce.10. Albanese, M. A. and S.

Albanese, M. A. and S. Mitchell, 1993, "Problem-Based Learning: A Review of Literature on its Outcomes and Implementation Issues," Academic Medicine, Vol. 68, pp. 52-81.11. Prince, M. J. and R. M. Felder, 2006, "Inductive Teaching and Learning Methods: Definitions, Comparisons, and Research Bases," Journal of Engineering Education, Vol. 95, No. 2, pp. 123-138.12. Dym, C. L., A. M. Agogino, O. Eriz, D. D. Frey and L. J. Leifer, 2005, "Engineering Design Thinking, Teaching and Learning," Journal of Engineering Education, 95, 1, 103-120.13. Davison, R. C., 2009, "Engineering Curricula: Understanding the Design Space and Exploring the Opportunities," National Academy of Engineering, Washington D.C. 14. Jamieson, L. and J. Lohmann, 2009, "Phase 1 Report: Creating a Culture for Scholarly and Systematic Innovation in Engineering Education," American Society of Engineering Education.15. Felder, R. M. and M. J.

and J. Lohmann, 2009, "Phase 1 Report: Creating a Culture for Scholarly and Systematic Innovation in Engineering Education," American Society of Engineering Education.15. Felder, R. M. and M. J. Prince, 2007, "The Case for Inductive Teaching," ASEE PRISM, 17, 2, 55.16. Prince, M. J. and M. Vigeant, 2006, "Using Inquiry-based Activities to Promote Understanding of Critical Engineering Concepts," American Society for Engineering Education Annual Conference, Chicago, IL.17. Hmelo-Silver, C. E., 2004, "Problem-based learning: What and How Do Students Learn?," Educational Psychology Review, 16, 3.18. Barrows, H. S., 1996, "Problem-based Learning in Medicine and Beyond: A Brief Overview," New Directions for Teaching and Learning, 68, 3-12.19. Ormrod, J. E., 2008, Human Learning (5th ed.), Pearson, Columbus, OH. 20. Dunlap, J.

Ormrod, J. E., 2008, Human Learning (5th ed.), Pearson, Columbus, OH. 20. Dunlap, J. C., 2005, "Problem-based Learning and Self-Efficacy: How a capstone course prepares students for a profession," Educational Technology Research and Development (ETR&D), 53, 1, 65-85.21. Bransford, J. D., A. L. Brown and R. R. Cocking, 2000, How People Learn: Brain, Mind, Experience and School, National Academy Press, Washington D.C.22. Blumenfeld, P. C., E. Soloway, R. W. Marx, J. S. Krajcik, M. Guzdial and A. Palincsar, 1991, "Motivating Project-based Learning: Sustaining the Doing, Supporting the Learning," Educational Psychologist, 26, 3/4, 369.23. Perrenet, J. C., P. A. J. Bouhuijs and J. G. M. M. Smits, 2000, "The Suitability of Problem- Based Learning for Engineering Education: Theory and Practice," Teaching in Higher Education, 5, 3, 345-358.24. Barron, B. J. S., D. L. Schwartz, N. J. Vye, M. Allison, A. Petrosino and L.

SIGN FOR ADDITIVE MANUFACTURING OF KINEMATIC PAIRSShrey Pareek*, Vaibhav Sharma*, and Rahul Rai*Design Analytics Research and Technology (DART) Lab, Department of Mechanical and Aerospace Engineering, University at Buffalo-SUNY, Buffalo-NY, 14260AbstractWhile additive manufacturing processes are better suited for fabrication of parts with complex geometries, they face serious challenges whilst fabricating parts that require relative motion with respect to each other. The primary challenge in additive manufacturing of mechanisms is preventing the mating parts from bonding with each other during the fabrication process. In this paper the authors investigate design and additive fabrication of kinematic pairs that can move relative to each other. The paper outlines fabrication of kinematic pairs based on optimal clearance value for three basic lower order kinematic pairs, viz. revolute pair, prismatic pair, and cylindrical pair.

The paper outlines fabrication of kinematic pairs based on optimal clearance value for three basic lower order kinematic pairs, viz. revolute pair, prismatic pair, and cylindrical pair. Using empirical testing functional relationships between extractive force and clearance, and between moment and clearance have been developed. These functional relationships can be used by users to fabricate kinematic pairs using FDM based 3D printing processes. The efficacy of the proposed approach is demonstrated on 3D printed kinematic pairs and experimental validation studies.Introductionthe unquestioned advantage over conventional manufacturing techniques when it comes to fabrication of parts with complex geometries. However, additive manufacturing (Fused Deposition Modeling in particular) encounters serious challenges whilst fabricating parts that require relative motion with respect to each other.

However, additive manufacturing (Fused Deposition Modeling in particular) encounters serious challenges whilst fabricating parts that require relative motion with respect to each other. Generally such models require that each part be produced separately and then assembled together. However, due to inherent inaccuracies of 3D printing these assemblies seldom function correctly. For instance, inaccurate circularity of a 3D printed shaft may prevent its smooth movement inside the hub. This problem can be eliminated through allocation of appropriate clearance. An appropriate clearance is one that allows for unobstructed functioning of the assembly while avoiding unwanted motions/vibrations such as wobbling in case of a shaft and hub.Additive manufacturing holdsCurrent 3-D printing technologies require that the user make amends to the original CAD model to account for appropriate clearance. However, this clearance is user defined.

However, this clearance is user defined. This leads to in proper sub-optimal clearance specification. Sub-optimal clearances cause problems functioning of the mechanism. Additionally, sub-optimal clearance specification necessitates multiple design iterations to identify optimal clearance in an ad-hoc manner. Multiple design iterations make the entire process entire process time consuming. The outlined method in this paper provides structured and effective way of determining optimal clearance for fabrication of 3D printed kinematic pairs. The suggested method is also fast, robust, and eliminates ad-hoc design iterations.Revolute and prismatic pairs have rotational and slidingrelationships w.r.t. each other respectively. A cylindrical pair exhibits both rotary and sliding relationships. The paper provides the optimal clearance value based on the user’s needs for three basic lower order kinematic pairs, viz. revolute pair, prismatic pair, and cylindrical pair (Figure 1).

The paper provides the optimal clearance value based on the user’s needs for three basic lower order kinematic pairs, viz. revolute pair, prismatic pair, and cylindrical pair (Figure 1). For the purpose of this paper we focus on clearance and interference fits only [1]. Tight fit is not considered as the proposed method allows the designers to use the 3D printed part directly i.e. no finishing operations are carried out on the 3D printed part and 3D printed kinematic parts can be used ‘as is’. For the purpose of this study a Fused Deposition Modeling (FDM) MakerBot Replicator 2X® is used as the 3D printing device.

For the purpose of this study a Fused Deposition Modeling (FDM) MakerBot Replicator 2X® is used as the 3D printing device. The results can be extended to other additive manufacturing techniques and 3D printing devices using suitable adjustments.Figure 1 - Schematic representation of Kinematic Pairs.Literature ReviewThe following section reviews previous work done in the field of additive manufacturing of assemblies and studies focused on understanding the relationship between surface roughness and limits, fits, and tolerances. Although the focus of the majority of the existing work is on interference fits, they provide us with vital information that can be extrapolated for clearance fits as well.Preventing the mating parts from bonding with each other during the fabrication process is a major concern in additive manufacturing.

For example, the CAD model of a piston-cylinder assembly ignores the clearance gap and the diameters of both the piston and cylinder is considered the same. This equal diameter is read as a single entity by the slicing algorithm, leading to fusion of the two parts. There are various methods implemented to counter this problem; Lipson et al. [2] successfully reproduced fully functional and preassembled kinematic models of Reuleaux’s kinematic pairs.Clearance between parts is responsible for determining the type of fit [2]. Studies have shown that in case of clearance fits, contact area increases with increase in load. Also the magnitude of circumferential stress depends on whether or not friction is considered [3]. Joint clearance is a critical deciding factor if the model is to exhibit the preferred movement.

Also the magnitude of circumferential stress depends on whether or not friction is considered [3]. Joint clearance is a critical deciding factor if the model is to exhibit the preferred movement. A lower clearance results in a tight fit restricting movement and difficulty in removing the support structure or powder from the gap and in some cases a water jet is used to remove the material [4]. To avoid tight fits Chen et al. suggested filleting or chamfering to ensure that there is enough contact area and no instability [5]. High clearance yields the joint unstable [6, 7]. Hence optimum joint clearance should be incorporated to get a functional joint. Chen et al. developed a drum pin joint by hit and trial method to minimize the joint clearance to the optimum level. This clearance also increased the joint strength simultaneously [6].The effect of clearance in joint stiffness has also been studied [3].

This clearance also increased the joint strength simultaneously [6].The effect of clearance in joint stiffness has also been studied [3]. It has also been shown that the fits in SLA depend heavily on surface roughness [8]. Yang et al. [9] study the effect of surface roughness on interference fit between a shaft and a hub. They compare the surface roughness with assembly pressure and fit. Additionally, they also outline a numerical formulation for loss of tightening. Using the numerical formulation they conduct experiments to show that surface texture greatly influences strength of fit. According to there numerical formulation extractive strength is a function of the friction coefficient and nominal contact pressure. The influence of surface finish on interference fits is further studied by Ramachandran et al. [10]. Although they study the load carrying capacity of fits w.r.t. the surface roughness, their approach and experimental investigations are of importance.

[10]. Although they study the load carrying capacity of fits w.r.t. the surface roughness, their approach and experimental investigations are of importance. The above observations serve as building blocks towards the formulation of the method proposed in this paper. Details related to these building blocks are outlined in the subsequent sections. It should be noted that the effect of stresses on the clearance are beyond the scope of this study.In case of Stereo Lithography (SLA), inserts of different materials are used [11]. Another technique is the use of a second release material placed in the gaps between the movable parts. This material is later etched, dissolved, melted or blown away [9]. Removal of support material might lead to a change in alignment of mating parts [9].

This material is later etched, dissolved, melted or blown away [9]. Removal of support material might lead to a change in alignment of mating parts [9]. In cases where no support structures are used instead of inserts, the clearance should be adjusted to ensure that the structure does not crash on removal of support structures [12].Further, most 3D models do not specify how their parts move or interact with each other. This leads to difficulty in visualization of the final part. For example a rolling shaft-hub pair may be treated as a sliding couple. N. Mitra et al. [12] developed a semi-automatic technique that determines the motion of parts and their causal relationships based on their geometry. This technique requires user assistance for successful analysis of CAD models. We propose a simpler technique that determines the correct clearance on the basis of conditions pertinent to each type of motion.

We propose a simpler technique that determines the correct clearance on the basis of conditions pertinent to each type of motion. It is assumed that parts rotate about a symmetric axis and/or translate along translation symmetry directions. These assumptions help in identification of translational or rotational axes and their degrees of freedom. If a part (P) remains unchanged under some transformation Tj i.e. Tj (P) = P, it is symmetric about that transformation axis [11].J Cali et al. [13] developed an automated technique that fits a 3D model into printable joints. They also focused on the aesthetic aspects of the final printed material. Adjustments were made in the model for it to exhibit internal friction. Our study ignores the aesthetic aspects and the internal friction of the model, and focuses mainly on functioning of the mechanisms.

Our study ignores the aesthetic aspects and the internal friction of the model, and focuses mainly on functioning of the mechanisms. Most of the aforementioned studies focus on use of SLA for additive manufacturing of mechanisms, whereas we use FDM as the premise for this study.MethodologyUsing computer simulations and experimental analysis we deduce functional relationships between clearance and extractive force, and clearance and moment required to rotate a rotational pair. These simulations and experiments are explained in the current section. During assembly all CAD packages require the designer to define the relative motion between components.

These simulations and experiments are explained in the current section. During assembly all CAD packages require the designer to define the relative motion between components. Based on the relations defined in the assembly model the empirically deducedfunctional relationships may be used to select the relevant conditions pertaining to each type of kinematic pair.Experimental	
 AnalysisExtractive	
 Force	
 MeasurementMoment	
 MeasurementExperimental	
 Results/	
 Functional	
 RelationshipsFinal	
 RelationshipsComputer	
 SimulationsExtractive	
 Force	
 MeasurementIdeal	
 Results/	
 RelationshipsFigure 2 - Schematic Representation of the proposed methodologyFigure 2 provides a basic process the overall methodology. study comprises of experimental computer and analysis simulations to study the effect of change in clearance on extractive force and moment required to rotate a rotational pair (revolute pair pair).

study comprises of experimental computer and analysis simulations to study the effect of change in clearance on extractive force and moment required to rotate a rotational pair (revolute pair pair). Extractive force is defined as the is minimum axial sufficient to displace a linear pair (prismatic pair and cylindrical pair). Results obtained from the and experimental computer (ideal results) are analyzed separately and w.r.t. each other to arrive at the final functional relationships.flow ofTheandcylindricalloadthatanalysis simulationsMoment Measurement HubWeightShaftsArmForce Measurement HubFigure 3 – Components used for ExperimentationExperimentUsing Acrylonitrile Butadiene Styrene (ABS) as and base material the MakerBot Replicator 2X® printer as FDM the is experimental carried out. To imitate the pairs three kinematic (prismatic, and cylindrical) a shaft and hub arrangement is used.

To imitate the pairs three kinematic (prismatic, and cylindrical) a shaft and hub arrangement is used. The diameter of the hub is kept constant at 20 mm and the is varied shaft diameter from 19.20 mm to 19.75 mm in increments of 0.05 mm. Diameters above 19.75 mm are ignored as they require chamfering operations of the end faces in order to insert the shaft into the hub. Since our process targets at using the 3D printed parts directly without any machining operations, samples with diameter greater than 19.75 mm are disregarded. The length of the shaft and hub are fixed at 50mm and 15mm respectively. An arm of length 47.5 mm is also provided at one end of the shafts in order to apply the moment (Figure 1). A total of 60 shaft samples (five samples for each diameter; 19.20 mm, 19.25 mm, 19.30 mm, 19.35 mm, 19.40 mm, 19.45 mm, 19.50 mm, 19.55 mm, 19.60 mm, 19.65 mm, 19.70 mm, and 19.75mm) are printed using standard settings (Table 1).

Rafts and support materials are used and the samples are printed vertically in order to maintain the circularity of the shaft and hub. Two different hub samples, one for force and the other for moment measurements are printed. The hub samples double up as fixtures for the experiment (Figure 3).Force Measurement HubMoment Measurement HubanalysisrevoluteShaftsFigure 4 – Disassembled Hubs and ShaftsTable 1 – Printing ParametersNozzle Movement Speed 150 mm/sPrint Speed90 mm/sNozzle Temperature220oCBed Temperature150oCLayer Height0.20 mmExtractive ForceIn order to calculate the extractive force required to achieve linear motion, the shaft is housed vertically in the force measurement hub and is loaded axially from the top (Figure 5). Extractive force for each diameter is taken as that minimum load which is sufficient to overcome the static friction between the hub and the shaft.

Extractive force for each diameter is taken as that minimum load which is sufficient to overcome the static friction between the hub and the shaft. Standard weights are used to obtain a roughFigure 5 – Extractive Force Measurementestimate of the load required and water filled beakers are used to obtain finer results. All samples are arranged in the same orientation while carrying out the experiment. As stated previously five samples for each diameter are used. Self-weight of each shaft is also considered in the analysis. All measurements are carried out in grams and are accurate up to two decimal places.Figure 6 – Moment MeasurementMomentThe moment required to overcome the static friction in the circumferential direction for each sample is measured. In this case the shaft is housed horizontally inside the moment hub and load is applied at the end of the arm (Figure 6).

In this case the shaft is housed horizontally inside the moment hub and load is applied at the end of the arm (Figure 6). The orientation of the samples is kept the same and standard weights and water filled bags are used for rough and fine load measurements respectively.Computer SimulationsIn order to gain an insight on the how the results obtained from the above experimental analysis differ from the ideal results, computer simulations of the above experiments were carried out using Finite Element Analysis (ANSYS®). Similar to the experimental setting, extractive force and moment variation w.r.t change in diameter of the shaft (clearance) are carried out. During FEA simulations, the dimension of the hub is kept at a constant diameter of 20mm and the diameter of the shaft is varied from 19 mm to 21mm over 100 data points. These data points are randomly generated by the software program.

These data points are randomly generated by the software program. For each value of the shaft diameter, the shaft is given a longitudinal displacement (0.1 mm) in case of the sliding and cylindrical pair, and an angular displacement (1o) in case of the revolute and cylindrical pair. Variation of extractive force/moment required on the shaft w.r.t. the varying diameter (clearance) is calculated and plotted. Figure 7 depicts the specimen used for the simulations. Simulations offorce and moments transmission across the shaft and hub are also carried out. Figure 8 demonstrates force transmission simulation for a 19.982 mm shaft.Experimental and Computational AnalysisExperimental AnalysisOnce the force and moment data points are obtained, they are averaged over the five samples at each diameter (clearance) value. These averaged points are then plotted against the clearance and polynomial regression analysis [14] is carried out in order to obtain the best fitting curve.

These averaged points are then plotted against the clearance and polynomial regression analysis [14] is carried out in order to obtain the best fitting curve. Figure 9 shows the graph of Extractive Force v/s Clearance for the experimented data points. The blue circles denote the averaged data points obtained by the experimental analysis and the black line is the best fitting cubic curve. A very high goodness-of-fit value (R2) of 0.9789 assures that the generated function is a good estimate of the relation between extractive force and clearance. Similarly Figure 10 shows the variation of moment as a function of clearance. The results are summarized in Table 2.Figure 7 - SpecimenComputer SimulationsFigures 11 and 12 show the plots the Extractive Force and generated for Moment computer using simulations. It should be noted here that 100 different values of clearance were used. Similar to the previous case curve fitting is carried are approximated.

This observation complies with the common knowledge about fits and their relationship with these forces. Careful observation of the results obtained from computer simulations (Figure 11-12) show that as the clearance is decreased the value of force (and moment) is constant at 0 N (N-mm) initially. A sudden increase in force and moment values is observed below clearances of 0.05 mm. These results suggest an optimal clearance within the range of 0 to 0.05 mm. However realization of a clearance of this small magnitude is not possible with the current 3D printer being employed. The computer simulations are thus deemed as ideal results and are overpowered by the experimental results, which are discussed in the next paragraph. Further, results obtained from force transmission simulations (Figure 8) suggest 100% force transmission and show no (zero) energy loss between the shaft and the hub.

Further, results obtained from force transmission simulations (Figure 8) suggest 100% force transmission and show no (zero) energy loss between the shaft and the hub. These results assert the differences between ideal (computer simulation) results and experimental results. There is always some loss in transmitted force (energy loss) in any kinematic pair. Energy loss is not captured accurately by computer simulations. The above observations augment the ineffectiveness of computer simulations and hence simulation results were discarded and only experimental results were used for deriving functional relationships.Graphs obtained from experimental analysis (Figures 9-10), are more useful for practical purposes. The difference in experimental and simulation results can be attributed to the0.3inaccuracies of 3D printing.

The difference in experimental and simulation results can be attributed to the0.3inaccuracies of 3D printing. Phenomenon such as warping or vibrations leads to inaccurate geometries (eccentricity of the shaft and hub in this case).ResultsThe final functional relationships are obtained using empirical analysis and are shown below. In the following equations, y denotes the Extractive Force/Moment (Equation 1 and 2) and x denotes the optimal clearance corresponding to that force/moment value.y = -715.38x3 + 1428.8x2 - 956.37x + 216.06 (1) (Extractive Force v/s Clearance)y = 1287.1x2 - 2479.2x + 1181.1 (2) (Moment v/s Clearance)Extractive force and moment conditions are the active conditions for a sliding pair and a revolute pair respectively. In case of a cylindrical pair both conditions are active. This pair is treated as a special case wherein the user selects one condition as the active one.

In case of a cylindrical pair both conditions are active. This pair is treated as a special case wherein the user selects one condition as the active one. For instance, if the user chooses the moment condition to be the active one, equation 2 is used to obtain the optimal clearance, which can then be used in equation 1 to obtain the corresponding extractive force.

The relationship between moment and extractive force pertaining to a cylindrical pair can be summarized using Figure 13 and Equation 3.y = 108.3x0.4627 (3) (Moment v/s Extractive Force)800Moment v/s Extractive Force700y	
 =	
 108.3x0.4627	
 R²	
 =	
 0.98886600)m m N500(t n e m o M40030020010000102030405060Extractive Force (N)Figure 13 – Moment v/s Extractive Force for Cylindrical Pair Design70ValidationIn order to show the utility of the proposed method, we outline the following examples.Example 1 – Prismatic PairAssuming for a particular prismatic pair (shaft-hub pair in this case), an extractive force of 30 N is required. In order to obtain the optimal clearance for this particular application, the Extractive Force – Clearance relationship (equation 1) is used.

In order to obtain the optimal clearance for this particular application, the Extractive Force – Clearance relationship (equation 1) is used. Using this relation the optimal clearance is determined as 0.331 mm.To validate the usefulness of this result, a shaft-hub pair with clearance 0.331 mm is 3D printed using the printing parameters listed in Table 1. It should be noted that for validation purpose, we used a different 3D printer but of the same make (MakerBot Replicator 2X ®). For the sake of consistency the hub is kept at a diameter of 20 mm and the shaft diameter is taken as 19.669 mm (clearance 0.331 mm). Extractive force measurements as described in Section 3.1.1 was carried out. The analysis yielded an extractive force value of 26.82 N as opposed to the initial assumption of 30N.

Extractive force measurements as described in Section 3.1.1 was carried out. The analysis yielded an extractive force value of 26.82 N as opposed to the initial assumption of 30N. This slight deviation in the observed value can be attributed to inherent inaccuracies of 3D printing process.Example 2 – Revolute PairAssuming a moment requirement of 200 N-mm and using equation 2 (Moment-Clearance Relation), the optimal clearance value is determined as 0.556 mm. Similar to example 1, this value is validated using a moment experiment on a 3D printed shaft-hub pair with clearance 0.556 mm (Section 3.2.2). The observed moment value was found to be 232.69 N-mm. This value differs from the original assumed value of 200 N-mm. This variation is acceptable and can be termed as the error of the proposed method.Example 3 – Cylindrical PairA cylindrical pair requires linear as well as rotary motion.

This variation is acceptable and can be termed as the error of the proposed method.Example 3 – Cylindrical PairA cylindrical pair requires linear as well as rotary motion. For this particular scenario we use the prismatic pair from Example 1 and add a rotary relation to it. As described in section 5, we select either the extractive force condition or the moment condition as the active one. In this case, selecting extractive force as the active condition with a value of 30 N, we determine the corresponding moment value as using equation 3 as 522.50 N-mm. Knowing the experimental extractive force value as 26.82 N from example 1, the moment value is measured using the method employed in example 2. The observed moment value for this case is 473.82 N-mm. Again a deviation from the calculated value (522.50 N-mm) is observed.As mentioned earlier the deviations observed in the above cases can be attributed to the inaccuracies and inconsistencies of 3D printing.

Since the measured values are averaged over the five samples at each diameter, the functional relationships are bound to be erroneous. Despite these inherent errors, the above examples demonstrate that our method has practical utility.Conclusion and Future WorkThe primary challenge in additive manufacturing of kinematic linkages is preventing the mating parts from bonding with each other during the fabrication process. This problem can be eliminated through allocation of appropriate clearance. The paper addresses the challenge of 3D printing of movable linkages and proposes a novel and robust technique that optimizes the clearance between these linkages. Using experimental analysis functional relationships between extractive force and clearance; and between moment and clearance was developed. The clearances obtained from the deduced functional relationships were shown to be optimal for clearance and interference fits.

The clearances obtained from the deduced functional relationships were shown to be optimal for clearance and interference fits. This optimal clearance assures proper functioning of the 3D printed joints and avoids unnecessary motion/vibration.The slight error in the method can be attributed to inherent inaccuracies of the 3D printing process and errors from experimental estimations (averaging values over five samples at each diameter/clearance). The presented approach could be used for most FDM printers and can be easily extended to other dimensions and additive manufacturing processes. For the purpose of this study, effects due to internal friction and stresses have been ignored. Incorporation of these factors into the study may lead to more accurate results.

For the purpose of this study, effects due to internal friction and stresses have been ignored. Incorporation of these factors into the study may lead to more accurate results. The study has focused on lower order kinematic pairs (sliding, prismatic and cylindrical) and needs to be extended to higher order linkages (ball and socket joint, universal joint) in order to enable 3D printing of variety of kinematic linkages. The study has focused on clearance as the sole design variable. Increasing the number of variables by varying the printing parameters (nozzle, velocity, bed temperature, and material) can make the method more robust. Studies exploring the relationship between surface roughness and extractive force/moment could be particularly beneficial for 3D printing parts with clearance fits.References[1] Mechanical engineer's reference book. Vol. 9. Butterworth Heinemann, 1998.[2] Lipson, Hod, Francis C. Moon, Jimmy Hai, and Carlo Paventi.

SIGN FOR ADDITIVE MANUFACTURING: A METHOD TO EXPLORE UNEXPLORED REGIONS OF THE DESIGN SPACEDavid W. Rosen The George W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332-0405 david.rosen@me.gatech.eduABSTRACTAdditive Manufacturing (AM) technologies enable the fabrication of parts and devices that are geometrically complex, have graded material compositions, and can be customized. To take advantage of these capabilities, it is important to assist designers in exploring unexplored regions of design spaces. We present a Design for Additive Manufacturing (DFAM) method that encompasses conceptual design, process selection, later design stages, and design for manufacturing. The method is based on the process-structure-property-behavior model that is common in the materials design literature. A prototype CAD system is presented that embodies the method.

The method is based on the process-structure-property-behavior model that is common in the materials design literature. A prototype CAD system is presented that embodies the method. Manufacturable ELements (MELs) are proposed as an intermediate representation for supporting the manufacturing related aspects of the method. Examples of cellular materials are used to illustrate the DFAM method.1INTRODUCTION1.1 Design for Additive ManufacturingDesign for manufacturing (DFM) has typically meant that designers should tailor their designs to eliminate manufacturing difficulties and minimize costs. However, the improvement of rapid prototyping, or Additive Manufacturing (AM), technologies provides an opportunity to re-think DFM to take advantage of the unique capabilities of these technologies. Several companies are now using AM technologies for production manufacturing.

Several companies are now using AM technologies for production manufacturing. For example, Siemens, Phonak, Widex, and the other hearing aid manufacturers use selective laser sintering (SLS) and stereolithography (SL) machines to produce hearing aid shells, Align Technology uses stereolithography to fabricate molds for producing clear braces (“aligners”), and Boeing and its suppliers use SLS to produce ducts and similar parts for F-18 fighter jets. In the first three cases, AM machines enable one-off, custom manufacturing of 10’s to 100’s of thousands of parts. In the last case, AM technology enables low volume manufacturing and, at least as importantly, piece part reductions to greatly simplify product assembly. More generally, the unique capabilities of AM technologies enable new opportunities for customization, improvements in product performance, multi-functionality, and lower overall manufacturing costs.

More generally, the unique capabilities of AM technologies enable new opportunities for customization, improvements in product performance, multi-functionality, and lower overall manufacturing costs. These unique capabilities include:Shape complexity: it is possible to build virtually any shape, lot sizes of one are practical, customized geometries are achieved readily, and shape optimization is enabled.Material complexity: material can be processed one point, or one layer, at a time, enabling the manufacture of parts with complex material compositions and designed property gradients.Hierarchical complexity: hierarchical multi-scale structures can be designed and fabricated from the microstructure through geometric mesostructure (sizes in the millimeter range) to the part-scale macrostructure.New CAD and DFM methods are needed in order to take advantage of these capabilities.

In the hearing aid and aligner cases, new CAD systems had to be developed to enable efficient shape modeling and part design. During a U.S. government sponsored study of European researcher groups [18], many researchers said that they foresaw the lack of capable CAD tools as a serious impediment for their research and for the utilization of AM technologies for production manufacturing applications. However, if suitable CAD and DFM methods and tools can be developed, designers can design devices with significantly improved performance that fully utilize material, and with efficient manufacturing processes. With the shape, material, and hierarchical complexity capabilities, DFM can move from an emphasis on cost minimization to a focus on achieving heretofore unrealizable capabilities. Hence, a new definition of DFM can be proposed.

Hence, a new definition of DFM can be proposed. DFM for Additive Manufacturing (DFAM) is the:Synthesis of shapes, sizes, geometric mesostructures, and material compositions and microstructures to best utilize manufacturing process capabilities to achieve desired performance and other life-cycle objectives.In order to achieve this new concept of DFAM and enable wide ranges of new applications, new approaches, methods, and tools are needed. The focus in this paper is on the application of cellular materials to replace bulk materials. Cellular materials provide many more design variables, but with those variables comes more complexity. New design, analysis, and manufacturing technologies are needed as a result.1.2 Cellular MaterialsThe concept of designed cellular materials is motivated by the desire to put material only where it is needed for a specific application.

From a mechanical engineering viewpoint, a key advantage offered by cellular materials is high strength accompanied by a relatively low mass. These materials can provide good energy absorption characteristics and good thermal and acoustic insulation properties as well [8]. Cellular materials include foams, honeycombs, lattices, and similar constructions. When the characteristic lengths of the cells are in the range of 0.1 to 10 mm, we refer to these materials as mesostructured materials. Mesostructured materials that are not produced using stochastic processes (e.g. foaming) are called designed cellular materials. In this paper, we focus on designed lattice materials.In the past 10 years, the area of lattice materials has received considerable attention due to their inherent advantages over foams in providing light, stiff, and strong materials [1]. Lattice structures tend to have geometry variations in three dimensions; some of our designs are shown in Figure 1.

Lattice structures tend to have geometry variations in three dimensions; some of our designs are shown in Figure 1. As pointed out in [5], the strength of foams scales as ρ1.5, whereas lattice structure strength scales as ρ, where ρ is the volumetric density of the material. As a result, lattices with a ρ = 0.1 are about 3 times stronger than a typical foam. The strength differences lie in the nature of material deformation: the foam is governed by cell wall bending, while lattice elements stretch and compress. The examples in Fig. 1 utilize the octet-truss (shown on the left), but many other lattice structures have been developed and studied (e.g., kagome, Kelvin foam). We have developed methods for designing lattice mesostructure for parts [25] and have developed design- for-manufacturing rules for their fabrication in SL.Methods of continuum mechanics have been applied to various mesostructured materials. Ashby and co-workers wrote a book on metal foam design and analysis [1].

Ashby and co-workers wrote a book on metal foam design and analysis [1]. They and others have applied similar methods to the analysis of lattice structures. The octet truss in Fig. 1 has been extensively analyzed. Deshpande et al. [5] treated the octet truss unit cell as a collection oftension-compression bars that are pin-jointed at vertices and derived analytical models of their collapse behavior for many combinations of stresses. Their results match finite element model behavior well, but tend to under-predict the strength and stiffness of octet trusses due to their assumption of pin-jointed vertices. Wang and McDowell [25] extended this study to include several other lattice cells. Recently, we have been developing a more general analytical model of lattice behavior [11]. From our general model, models for octet and other lattice structures can be derived. We base our model on a single vertex with a collection of struts incident on that vertex.

From our general model, models for octet and other lattice structures can be derived. We base our model on a single vertex with a collection of struts incident on that vertex. This vertex model will be our base “unit cell” for representation and modeling purposes.wduvFigure 1. Octet-truss unit cell and example parts with octet-truss mesostructures.1.3 Requirements for DFAMThe concept of mesostructured materials is motivated by the desire to put material only where it is needed for a specific application. Achieving high stiffness or strength and minimal weight are typical objectives. Multifunctional requirements are also common, such as structural strength and vibration absorption.

Achieving high stiffness or strength and minimal weight are typical objectives. Multifunctional requirements are also common, such as structural strength and vibration absorption. The area of compliant mechanisms shares the same motivation, where the local compliance of the structure enables the mechanism to perform specified motions.We hypothesize that designed mesostructures will enable structures and mechanisms to be designed that perform better than parts with bulk or non-designed mesostructures, particularly for multifunctional applications. Testing this hypothesis requires the ability to bridge the meso to macro size scales. To do this, we need to first recognize some requirements on DFAM methods and CAD-DFAM tools, with a related objective of utilizing the unique capabilities of AM technologies. The requirements that we propose include the capability to:Define and explore large, complex design spaces.

The requirements that we propose include the capability to:Define and explore large, complex design spaces. • Represent and design with hundreds of thousands of shape elements, enabling large complex design problems as well as designed material mesostructures.Represent complex material compositions and ensure that they are physically meaningful.Determine mechanical properties from material compositions and mesostructures across length scales.Ensure that specified shapes, material structures, and properties are manufacturable.In order to achieve these requirements, several new technologies are required.

Our approachhas five main elements (requirements that they address are appended):Process-Structure-Property-Behavior-Function framework for Design for Additive Manufacturing.Cellular materials: methods for topology layout, high-fidelity analysis, simplified analysis.“Manufacturable Elements” (MEL’s) that contain geometry, material, properties, and uncertainties in these quantities.DFAM templates that enable formulation and solution of typical DFM problems. • Improved search algorithms needed to explore large, complex design space.Each of these five new technologies is addressed, at least in part, in this paper. In the next section, the framework for our DFAM approach is presented, providing the larger context for this research. In Section 3, the technologies being developed for DFAM are presented, covering our bio-inspired conceptual design, cellular materials layout, and Manufacturing Elements methods.

In Section 3, the technologies being developed for DFAM are presented, covering our bio-inspired conceptual design, cellular materials layout, and Manufacturing Elements methods. An example is presented in Section 4 that covers part design, manufacturability analysis, process planning, and process simulation. Conclusions are drawn in the final section.2 FRAMEWORK FOR DFAMWe will borrow the process-structure-property relationships framework from the materials science field to model a design [13]. The manufacturing process space, P, consists of process plans with sequences of operations and values of process variables. Structure space, S, contains information about the geometric, topological, and material structures of a design. Property space T contains information about part properties that are derivable from S using physical principles; e.g., mechanical, thermal, and electrical properties.

Property space T contains information about part properties that are derivable from S using physical principles; e.g., mechanical, thermal, and electrical properties. Behavior space, B, contains information about a part’s actual behavior given some loading and boundary conditions, while function space, F, describes the desired behavior of the system. The relationships among these spaces are shown in Fig. 2.Mappings are defined among these spaces. Mapping Φ represents a manufacturing analysis that determines material composition and microstructure, and possibly as-manufactured part shape from a process plan. Mapping Γ represents a material science analysis of a material, and possibly part geometry, to arrive at a set of mechanical and other properties. It is possible that this mapping can be determined once, then reused for different applications (not design- or part- specific).

2: Process-Structure-Property-Behavior Mappings.The mappings shown in Fig. 2 capture important relationships among design attributes across several size scales. However, additional richness enables the multi-scale aspects of both geometric and material models to be captured. The process, structure, and property models will be divided into geometric and material models in order to emphasize their different decompositions. Fig. 3 shows the framework from Fig. 2 with separate Material and Geometry levels. As one moves from right to left, the relevant size scales decrease. Similarly, when one moves from top to bottom, smaller size scales become important.

2 with separate Material and Geometry levels. As one moves from right to left, the relevant size scales decrease. Similarly, when one moves from top to bottom, smaller size scales become important. These levels and size scales will be further explained in the next section.PGPG PGSG SG SG SG SG SGTGTG TGB Behavior B Behavior B BehaviorGeometry GeometryΦ ΦΓ Γ∆ ∆Decreasing size scale Decreasing size scaleMaterial MaterialPM PM PM PM Process Process Process ProcessSM SM SM SM SM SM Structure Structure Structure Structure Structure StructureTM TM Property PropertyDecreasing size scale Decreasing size scale Fig. 3: Expanded DFAM Framework with Geometry and Material Layers.3 DFAM METHOD, SYSTEM, AND TECHNOLOGIES3.1 Bio-inspired Conceptual DesignThe overall DFAM method begins with a conceptual design stage that is based on biomimicry. Our approach is an extension of the Pahl and Beitz design method [15].

Our approach is an extension of the Pahl and Beitz design method [15]. For some key subfunctions identified by the designer, a bio-inspired approach can be used to leverage “solutions” from nature. Our method is called “reverse engineering biological systems” and is intended to help designers to develop solution and working principles by abstracting from the working principles used in biological system.

Resulting “biological strategies” can be used as creative stimuli in the search for engineering principles.Systematic Design Systematic Design Systematic Design Methodology Methodology MethodologyBiological Biological Biological System System System Identification Identification IdentificationPlanning and Planning and Planning and Clarification of Task Clarification of Task Clarification of TaskConceptual Conceptual Conceptual Design Design DesignAs seen in Figure 1, there are four key research the method for reverse engineering in areas biological systems identification, biological representation, biological strategy extraction, and strategy abstraction.

The uniqueness of this method lies in the last three research system identification is currently being addressed by other researchers [3,4,21,22].systems:biologicalareas, whereasbiologicalg g n n i i r r e e e e n n g g n n E E e e s s r r e e v v e e R Ri iBiological Biological Biological Representation Representation RepresentationBiological Biological Biological Strategy Strategy Strategy Extraction Extraction ExtractionStrategy Strategy Strategy Abstraction Abstraction AbstractionAbstract essential Abstract essential Abstract essential Abstract essential Abstract essential Abstract essential problem problem problem problem problem problemConceptual Conceptual Conceptual Establish function Establish function Establish function Establish function Establish function Establish function Systematic Design Systematic Design Systematic Design structures structures structures Design Design Design structures structures structures Methodology Methodology MethodologySearch for working Search for working Search for working Search for working Search for working Search for working principles principles principles principles principles principles…….…….

…….……. …….…….A key step in extracting biological strategies is to develop a model of the biological system’s behavior. Such behavior models are critical in identifying engineering systems where the biological system may be applied, as well as in adapting the biological system to the engineering system. We areEmbodiment Design Embodiment Design Embodiment DesignDetail Design Detail Design Detail DesignFig. 4 Conceptual design method.using hierarchical Petri nets [28] as the basis of our behavioral models. Specifically, we believe that modeling the discrete physical states of the biological system as places and changes in these discrete states as transitions in the Petri net framework, while still holding the properties of reachability, liveness, and boundedness.We have applied the bio-inspired conceptual design method to the design of morphing aircraft wing skins (based on the sea cucumber and human muscle) [27] and of artificial kidneys (based on human kidneys).

We will conclude this presentation by stating that the output of conceptual design should be a detailed behavior model as well as working principles for each function in the behavior model. This information drives later design stages. Since the method is not directly related to SFF, we will not explain it further in this paper.3.2 DFAM Method and SystemThe overall DFAM method consists of a traversal of the frameworks in Figs. 2 and 3 from Function to Process, then back again to Behavior. The traversal from Function to Process can be called design, where functional requirements are mapped to properties and geometry that satisfy those requirements to structures and through process planning to arrive at a potential manufacturing process. Reverse direction, one can simulate the designed device and its manufacturing process to determine how well it satisfies the original requirements.Fig. 5 shows the proposed DFAM system that embodies the method outlined above. To the right in Fig.

5 shows the proposed DFAM system that embodies the method outlined above. To the right in Fig. 5, the designer can define the DFAM synthesis problem, using an existing problem template if desired. For different problem types, different solution methods and algorithms will be available. Analysis codes, including FEA, boundary element, and specialty codes, will be integrated to determine design behavior. In the middle of Fig. 5, the heterogeneous solid modeler (HSM) is illustrated (heterogeneous denotes that material and other property information will be modeled). Libraries of materials and mesostructures enable rapid construction of design models. To the left, the manufacturing modules are shown. Both process planning and simulation modules will be included. After planning a manufacturing process, the idea is that the process will be simulated on the current design to determine the as-manufactured shapes, sizes, mesostructures, and microstructures.

After planning a manufacturing process, the idea is that the process will be simulated on the current design to determine the as-manufactured shapes, sizes, mesostructures, and microstructures. The as-manufactured model will then by analyzed to determine whether or not it actually meets design objectives.3.3 Synthesis MethodsTo date, we have used a synthesis method based on Particle Swarm Optimization (PSO), which is an extension of genetic algorithms (GA), to perform parametric and limited topological optimization of structures and compliant mechanisms. PSO simulates the movement of birds in a flock, where individuals adjust their flying according to their experience and other individuals’ experiences during searches for food [12]. It combines local search with global search, and enables cooperative behavior among individuals (“birds”), as well as the competition modeled using GA.

It combines local search with global search, and enables cooperative behavior among individuals (“birds”), as well as the competition modeled using GA. Hence, PSO often converges more quickly than GA and was selected for the design synthesis of cellular structures [7].In the future, we intend to adopt a two-stage method for multifunctional topology design applications that demand not only targeted structural performance but also satisfactory performance in a distinct secondary functional domain [18]. Intended secondary domains, such as conjugate heat transfer or vibration absorption, are governed by non-local, scale-dependent phenomena that are not directly amenable to standard homogenization or interpolation techniques underlying discrete or continuum topology optimization techniques.

For this class ofapplications, conventional approaches involve either selecting a standard topology [22] or identifying a final topology via conventional structural topology optimization [2,24] and thereby fixing its topology for subsequent multifunctional customization. Instead, a two-stage approach for multifunctional topology design is promising in which both topology and dimensions are adjusted for multifunctional performance requirements. For the first stage, a robust structural topology design process has been developed for designing a preliminary topology with structural performance that meets targets as closely as possible while remaining relatively insensitive to bounded adjustments in the topology itself and its dimensions. In the second stage, the topology is modified, within the acceptable bounds, to improve its multifunctional performance in a secondary domain.

In the second stage, the topology is modified, within the acceptable bounds, to improve its multifunctional performance in a secondary domain. The method relies on approximate physics-based models to facilitate rapid exploration of a broad design space and identification of promising multifunctional solutions that are verified subsequently with more detailed models.3.4 CAD = Structure + PropertyThe CAD system proposed here consists of the Structure and Property elements of Figs. 2 and 3. Our proposed geometric representation is a combination of implicit, non-manifold, and parametric modeling, with the capability of generating BReps when needed. Implicit modeling is used to represent overall part geometry, while non-manifold modeling is used to represent shape skeletons.

5: DFAM System and Overall Method.Design problems are formulated and solved within the Design box. To support geometric and structure/property reasoning needs, the CAD system maintains a high fidelity models of device geometry, material composition, and property distribution within the device. Process planning and manufacturing process simulations are supported within the Manufacturing module. The set of databases at the bottom of Fig. 5 illustrate the libraries of models and templates that are integrated in the system. Unit cells for cellular materials, properties of metals and polymers, models of manufacturing processes, and design problem formulations (templates) are among the integrated information.The approach taken to design with cellular materials is illustrated in Fig. 6. Typically, solid sections or thick walls of a part are to be replaced with a cellular structure in order to lighten the part, stiffen it, or for other functional reasons.

6. Typically, solid sections or thick walls of a part are to be replaced with a cellular structure in order to lighten the part, stiffen it, or for other functional reasons. The boundary surface of the part CAD model is partitioned into surface patches. Bezier or b-spline surface patches are fit to these patches. Amapped meshing approach [24] is used to fill the solid section or line the boundary with cellular structure. A solid model of the cellular material is then generated using our hybrid modeling approach [23]. The resulting solid cellular model is suitable for process planning (Section 3.5).Within the CAD system, non-manifold modeling is used to represent shape skeletons. For cellular structures, it is often sufficient to represent struts as line segments terminated by nodes. Radius parameters are associated with struts and nodes to enable reasoning about the 3D geometry of lattices and enable generation of 3D solid models, analysis models, and manufacturing models.

Radius parameters are associated with struts and nodes to enable reasoning about the 3D geometry of lattices and enable generation of 3D solid models, analysis models, and manufacturing models. We use a simple non-manifold model based on that of Gursoz et al. [9], which is particularly useful when representing lattices with skins.Divide bounding surface Divide bounding surface into surface patches into surface patchesSelect cell type Select cell type from library from libraryFill part volume with Fill part volume with conformal cells conformal cells0.5 0.51 1Upper Side Surface Upper Side Surface0.5 0.5Bottom Bottom0 00.5 0.5Lower Side Surface Lower Side SurfaceConvert topological model Convert topological model to solid model to solid modelFig. 6 Cellular modeling approach.3.5 Manufacturable ElementsA Manufacturable ELement (MEL) is a predefined, parameterized decomposition of a volumetric region of a part.

6 Cellular modeling approach.3.5 Manufacturable ElementsA Manufacturable ELement (MEL) is a predefined, parameterized decomposition of a volumetric region of a part. For the lattice structures under investigation in this paper, MEL definition is straightforward: scan patterns and scan variables are associated with each strut in a unit cell. Consider the octet unit cell in Fig. 1 and assume it is being built in a SL machine vertically upward. For each layer in the SL build process, the unit cell is sliced by a plane. For the vertical struts, the intersection of the plane and the strut is a circle. For slanted struts, the intersection is an ellipse, while for horizontal struts, the intersection is a rectangle. Each case can be handled readily.The cases for vertical and slanted struts are shown in Fig. 7.

Each case can be handled readily.The cases for vertical and slanted struts are shown in Fig. 7. The notation is as follows: r = strut radius, W0 = laser beam radius, θ = strut angle, rl = major axis of ellipse (with minor axis = r), and p = (px, py) = center of intersected circle or ellipse. The specific parameters in the cases were determined empirically and give reasonable results for typical SL resins and laser scanning speeds. For example, case a) rl ≤ 1.4 W0 or 1.6 W0 or other multiple of W0 could have been selected and the laser irradiation time can be determined easily. However, long irradiation times can cause cured struts to become too thick, while short irradiation times may not enable the layerto adhere to the previous one. The multiple 1.5 times W0 is a reasonable compromise value.

The multiple 1.5 times W0 is a reasonable compromise value. For horizontal struts, cases b), c), and d) apply.r rW0 W0W0 W0rl rlxm -xmxm xmScan is just a point1 scan vector from x = -xm to x = xm.xm = rl - W0Case a) rl ≤ 1.5 W0Case b) r ≤ 2 W0xm -xmScan 3 Scan 3 ym ymr rxm xmr rW0 W0Scan 1 Scan 1ym -ymScan 2 Scan 23 scan vectors: Scan 1 is same as case b),Scan 2 is at y = -ym, Scan 3 is at y = ym.x m=r l1−2 y m2 x W − 0 W0ym = rl - 0.8Case c) r ≤ 3 W0Case d) otherwiseFig. 7 Scan pattern cases for sliced struts.Using the standard SL exposure model from Jacobs [10], the irradiation time for points and scan speeds for lines can be computed easily.

7 Scan pattern cases for sliced struts.Using the standard SL exposure model from Jacobs [10], the irradiation time for points and scan speeds for lines can be computed easily. For reasonably long scan vectors (more than ~3 times the laser beam diameter), the exposure received at a point (y, z) in the vat by a scan along the x axis is given byE x y z ( , , )=2 πP L W V 0 se2 −2 2 y W 0e−z Dp(1)where PL = laser power [mW], Vs = scan speed [mm/s], and Dp = depth of penetration [mm] (taken to be a constant measure of a resin’s sensitivity to laser energy). SL resins are assumed to be cured (form a solid) when they receive exposure that is equal to or greater than a certain amount, called the resin’s critical exposure, Ec [mJ/mm2]. For a layer thickness of l = 0.1 mm, it is typical to cure the resin to a depth of 0.15 mm, or 1.5 times the layer thickness. This curedepth reaches a maximum along a scan’s centerline (when y = 0). Eqn.

This curedepth reaches a maximum along a scan’s centerline (when y = 0). Eqn. 1 can be rearranged to solve for Vs as follows in order to compute the scan velocity to give a cure depth of 1.5 times the layer thickness:V s=2 πP L W E 0 ce −  1.5 Dlp   (2)Substituting reasonable values for a SLA-250/50 machine (PL = 30 mW, W0 = 0.125 mm, Ec = 0.12 mJ/mm2, l = 0.1mm, Dp = 0.1524 mm) yields a scan speed of about 600 mm/s. At this speed, the width of a cured scan line is 0.172 mm for the numbers in this example, or about 2/3 of the laser beam diameter. The cure model presented briefly here has been implemented into a MEL for lattice unit cells fabricated using SL.

The cure model presented briefly here has been implemented into a MEL for lattice unit cells fabricated using SL. By adjusting scan speeds, it is possible to fine- tune a process plan such that lattice struts have appropriate sizes, which has been formulated as a parameter estimation problem and solved using nonlinear least-squares methods [19].3.6 Process PlanningProcess planning is denoted by the mapping Φ : S → P. Using the notation in Fig. 3, process planning consists of two parts, one dealing with geometry decomposition (Φ : SG → PG) and the other for assigning values to process variables (Φ : SM → PM) to process the material appropriately. Geometric decomposition and process modeling are governed by the specific type of MEL selected.

Geometric decomposition and process modeling are governed by the specific type of MEL selected. In this section, we will briefly present the process planning formulation.Parameter estimation, or “inverse design,” methods can be applied to AM process planning to enable plans to be designed that meet design requirements on shape, surface finish, tolerances, and potentially other properties such as stiffness. Inverse design methods were developed in the heat transfer area [14]. A typical application of inverse design methods is to layout heater elements in a furnace, where heater positions are to be adjusted to achieve a desired temperature distribution. Parameter estimation methods for SL will be used to achieve a desired surface finish. The surfaces of a part fabricated in SL are defined by where the resin reaches a high enough crosslink density to remain solid, which is related to the exposure received from the laser [10].

The surfaces of a part fabricated in SL are defined by where the resin reaches a high enough crosslink density to remain solid, which is related to the exposure received from the laser [10]. The challenge is to determine appropriate exposure levels for each laser scan as it draws part cross sections such that part surfaces are precisely positioned and shaped.The inverse design problem for SL process planning can be stated as: find exposure valuesalong each scan vector to minimize the deviation of exposure across part surfaces from the desired constant Ec value. A general problem mathematical formulation is shown in Fig. 8 [19]. The constraint models the height of the part at a point P on its surface, which has to be equal to the summation of the layer thicknesses at point P and the thickness of the Compensation Zone at point P.

The objective to be minimized models the deviation of exposure at a set of grid points (on the part’s down-facing surfaces) from the critical exposure, Ec.functionGiven: Geometry of the part g(x,y,z)Material properties: energy absorption, α, critical exposure, Ec. Find: LTi, OCip, CZp Satisfy:nConstraints:hp=+∑LT CZipi 1 = Bounds: on variablesMinimize:n∑( α eLT OC +iip)ei  α − ∑  m 1 =LT CZ + mp  −11 =for all grid points PFig. 8 Math formulation of general process planning problem.Since there are many more scan vectors than measurement points mj, least-squares solution techniques are appropriate. We can take advantage of the MEL model by utilizing MEL parameterizations to sample each MEL and to compute exposure values. The least-squares fitting problem can be formulated as follows. The squared error term is the square of the objective function from Fig. 8, denoted by D (Eqn. 3).

The least-squares fitting problem can be formulated as follows. The squared error term is the square of the objective function from Fig. 8, denoted by D (Eqn. 3). This error term is to be minimized, so the derivative of D, with respect to the vector of variables U, involves the Jacobian of the system. Since J is nonlinear, an iterative solution technique must be used to solve for the unknowns, which are the scanning velocities and some scan vector positions. Both Gauss-Newton and Levenburg-Marquardt methods [16] are frequently used to solve such problems. In our work, we use Matlab’s non-linear least-squares solver, lsqnonlin, which selects from Gauss-Newton and Levenburg-Marquardt algorithms to solve problems.4 EXAMPLEAs an example, a cover plate for an aerospace structure will be redesigned to use lattice structure to stiffen it. The cover plate is shown in Fig. 7. It is approximately 300x350 mm in size and 3 mm thick.

The cover plate is shown in Fig. 7. It is approximately 300x350 mm in size and 3 mm thick. The thickness will be increased to 9 mm to accommodate the lattice structure, while the skin thickness will be decreased to 1.5 mm. A typical design-manufacture scenario will be presented that includes the decomposition of the cover plate geometry into cells, the synthesis of the resulting cellular structure to achieve a desired stiffness with minimum weight, and the decomposition of the into manufacturing synthesized operations using MELs. The central region of the plate, inside of the bolt hole pattern, is offset by the desired thickness of the lattice structure. This central region is then decomposed into lattice cells by mapping one layer of octet truss cells into the region. A nominal size of 8x8x8 mm is chosen for the cells, which results in 14,960 struts.Fig.

A nominal size of 8x8x8 mm is chosen for the cells, which results in 14,960 struts.Fig. 9 Cover plate example part.geometryTo achieve the objectives of a target stiffness (modeled by a target deflection) and minimum weight, a shape optimization problem is solved. Lattice strut diameters are the design variables. Rather than allowing each of the strut diameters to be a variable, we adopt the strategy of grouping struts into clusters based on an initial structural analysis. Ten clusters were used, corresponding to 10 design variables for optimization, since in our experience good results are often achieved. Diameter variables can vary between 0.2 and 1.2 mm, corresponding to the minimum manufacturable strut size on the lower end. A reasonably large strut size is chosen for the maximum of the range; if the strut diameters become larger, the cells start to lose porosity.

A reasonably large strut size is chosen for the maximum of the range; if the strut diameters become larger, the cells start to lose porosity. The loading condition for size optimization is an area load in the plate center of 0.064 N/mm2 applied to a 60x60 mm area.Size optimization is performed in ANSYS. Input files are automatically generated for ANSYS from the cellular model. The first-order gradient optimization method is used. Results are shown in Tab. 1. The number of optimization iterations, maximum stress, maximum deflection of the cover plate, and part volume are reported, along with the values of the diameter variables. This problem was not sensitive to the number of clusters used. Convergence of thecluster diameters was smooth, with the lower stress clusters becoming thicker, some mid-level clusters fluctuating up or down, as shown in Tab. 1. The final cover plate design is shown in Fig. 8.

1. The final cover plate design is shown in Fig. 8. Note that various strut diameters can be seen in the zoomed view.Diameters D1 D2 D3 D4 D5 D6 D7 D8 D9 D10Initial [mm] 0.2 0.4 0.6 0.6 0.8 0.8 1.0 1.0 1.2 1.2Final [mm] 0.2 0.6 0.79 0.73 0.84 0.73 0.88 1.04 1.2 1.2Max. StressVolumeInitial [N/mm2] -278.15Initial [mm3] 8008Final [N/mm2] 57.33Final [mm3] 20,314Tab. 1: Lattice optimization results.The next step in the DFAM process is process planning to ensure manufacturability. Each unit cell of the lattice structure is represented by a lattice MEL from Section 3.4. Recall that the relative orientation of each strut to the build direction dictates how it will be decomposed into manufacturing operations, which, in the case of SL or SLS are laser scans. Each strut is modeled as a MEL to facilitate process planning and process simulation.

Each strut is modeled as a MEL to facilitate process planning and process simulation. Process planning results for two struts were presented in an earlier paper [17].Fig 10 Cover plate with optimized lattice structure (shown on only half of the plate).5 CONCLUSIONSTwo advances are reported in this paper. First, a new Design for Additive Manufacturing (DFAM) method was presented that supports part and specification modeling, bio-inspired conceptual design, cellular material layout, process planning, and manufacturing simulations. Second, Manufacturable ELements (MELs) were proposed as an intermediate representation for supporting the manufacturing related aspects of the method.

Second, Manufacturable ELements (MELs) were proposed as an intermediate representation for supporting the manufacturing related aspects of the method. Specific conclusions from this work include:Cellular material types provide one method for providing mesostructure within a part for achieving improved stiffness, strength, or other functional requirements, as compared to monolithic materials.The process-structure-property-behavior framework can be expanded for the design for additive manufacturing of structures and devices.Bio-inspired conceptual design is a promising “front end” for the DFAM CAD system presented here.Manufacturing ELements (MELs) enable process planning within discrete regions of a part and also enable process simulation.

MELs represent one approach to achievingshape-specific process planning, rather than a reliance on general purpose process planning methods, typical of the AM industry at present.Design for Additive Manufacturing should be concerned with the exploration of expanded design spaces, rather than the focus on constraints imposed by the manufacturing processes, as is typical of DFM methods.6 ACKNOWLEDGEMENTSWe gratefully acknowledge the U.S. National Science Foundation, through grants IIS- 0120663 and DMI-0522382, and the support of the Georgia Tech Rapid Prototyping and Manufacturing Institute member companies, particularly Pratt & Whitney, for sponsoring this work. The assistance of Jamal Wilson is gratefully acknowledged for Section 3.1.7 REFERENCES [1] Ashby, M.F.; Evans, A.; Fleck, N. A.; Gibson, L. J.; Hutchinson, J. W.; and Wadley, H. N. G.(2000) Metal Foams: A Design Guide, Butterworth-Heinemann, Woburn, MA.[2] Bendsoe, M.P., Kikuchi, N.

SIGN FOR ADDITIVE MANUFACTURING: AN INVESTIGATION OF KEY MANUFACTURING CONSIDERATIONS IN MULTI-MATERIAL POLYJET 3D PRINTINGNicholas A. Meisel and Christopher B. Williams Design, Research, and Education for Additive Manufacturing Systems Laboratory, Virginia TechABSTRACTThe PolyJet material jetting process is uniquely qualified to create complex, multi-material structures. However, there is currently a lack of understanding and characterization regarding important manufacturing considerations to guide designers in their use of the PolyJet process. This paper investigates key considerations necessary to ensure that proposed designs are manufacturable and that part properties are appropriate for the intended use.

This paper investigates key considerations necessary to ensure that proposed designs are manufacturable and that part properties are appropriate for the intended use. Considerations included in this paper include 1) minimum manufacturable feature size, 2) removal of support material from channels, 3) survivability of small features during water jet cleaning, and 4) the maximum self-supporting angle of printed parts in the absence of support material. The result of this work is an understanding of which geometric and process variables affect these manufacturing considerations. This understanding is crucial for the creation of a set of Design for Additive Manufacturing (DfAM) guidelines to help designers create ideal, manufacturable parts with less iteration and provide constraints for insertion into automated design processes such as topology optimization.Keywords: Design for Additive Manufacturing (DfAM), PolyJet, 3D Printing, Multiple Materials,Material Jetting1.

DESIGN FOR ADDITIVE MANUFACTURING (DfAM)Additive Manufacturing (AM) allows for the creation of parts of seemingly infinite complexity. While “complexity is free” for AM systems relative to traditional manufacturing technologies, there are still manufacturing constraints that must be considered when designing parts for AM. These manufacturing considerations vary between the different AM process types, with certain considerations playing a larger role in one process than in another. As an example, material extrusion processes must consider the significant material anisotropy that occurs in the layering direction, while direct metal processes must account for added support scaffolding and anchoring. As these constraints differ from traditional manufacturing processes, dedicated “Design for Additive Manufacturing” (DfAM) guidelines are needed to assist engineers in realizing products that make effective use of AM’s capabilities.

A crucial component of these DfAM guidelines is identifying which geometric and process parameters will affect the manufacturing constraints (e.g., does layer thickness affect anisotropy in material extrusion parts?).Modern automated design methods, such as topology optimization, can help engineers design parts that may have been almost unthinkable prior to the advent of AM. In fact, numerous researchers are investigating the use of topology optimization in the field of AM (see, for example [1–3]). An understanding of manufacturing considerations is necessary to ensure that these automated methods produce meaningful designs that can be accurately printed. At a basic level, this could be as simple as knowing the dimensions of the printer’s build volume to ensure that the algorithm cannot design a part that is larger than the machine can build.

At a basic level, this could be as simple as knowing the dimensions of the printer’s build volume to ensure that the algorithm cannot design a part that is larger than the machine can build. A more complex example is the incorporation of the anisotropic properties of printed materials in the analysis phase of topology optimization.1.1. Establishment of a DfAM FrameworkSeveral researchers are working to establish Design for Additive Manufacturing (DfAM) frameworks that will help designers to better account for the advantages and limitations of AM processes in theirdesigns. Rosen proposed a modified Process-Structure-Property-Behavior framework with a varying size scale [4]. The framework connects the behavior of parts back to key aspects of the manufacturing process. He utilized this framework to create a DfAM system for the creation of complex cellular structures.

The framework connects the behavior of parts back to key aspects of the manufacturing process. He utilized this framework to create a DfAM system for the creation of complex cellular structures. This framework is advantageous because the behavior, properties, and structure of a part can all be mapped back to the advantages and constraints of the manufacturing process used to produce it.Hascoet and co-authors proposed what they termed a “global approach” to the incorporation of DfAM into printed parts [5,6]. Within this framework they suggest combining the functional requirements of the finished part with the manufacturing constraints of an AM process to develop a final geometry; this approach can be likened to the backward mapping that might occur from behavior to process in Rosen’s framework. The global approach is in contrast to the “partial approach” where a part’s geometry is designed beforehand and then adjusted to ensure manufacturability with AM.

The global approach is in contrast to the “partial approach” where a part’s geometry is designed beforehand and then adjusted to ensure manufacturability with AM. Hascoet cites work by Harzheim and Graf as a suitable example of this global approach. Harzheim and Graf combined the concept of topology optimization with the limitations of traditional metal casting (rather than AM) to create final parts directly from functional specifications and manufacturing constraints [7,8].Maidin’s approach to DfAM relies more on detailed investigation of case studies than the more predictive analytic methods of Rosen and Hascoet. Maidin established a design feature database meant to spark inspiration in the designers hoping to use AM for creation of their products [9].

Maidin established a design feature database meant to spark inspiration in the designers hoping to use AM for creation of their products [9]. He looked to existing products from AM and attempted to distill their novel aspects into a taxonomy useable by those unfamiliar with AM (example features included “dual material products,” “internal structuring,” and “weight saving features”). Finally, Seepersad and coauthors used a similar experimental case study approach too establishing DfAM rules, but focused more on establishing clear quantitative inputs for DfAM. They performed a series of experiments to establish a designer’s guide for creation of holes, thin walls, text, and mating gears in the SLS process [10]. They created matrices to help provide guidelines for designers in common situations, e.g., a determination of how large embossed lettering should be in order to be legible when made via laser-sintering.1.2.

ContextAn understanding of process constraints will be of benefit not only for assisting designers in preparing parts, but also in the creation of automated design tools that consistently provide manufacturable results. As already mentioned, topology optimization (TO) is one such tool. In essence, TO is an automated design methodology where a general design domain is defined (with associated forces, supports, and responses). Material is then systematically distributed in the space, according to a particular algorithm. The final result is a part that efficiently and effectively uses material to satisfy all constraints and maximize or minimize an objective function. More complex algorithms are able to incorporate various manufacturing considerations during the optimization process (see, for example, [11,12]).

More complex algorithms are able to incorporate various manufacturing considerations during the optimization process (see, for example, [11,12]). The goal of the research in this paper is to better understand the AM considerations that would need to be accounted for in such a manufacturing-oriented algorithm.Due to its unique capability to produce multi-material parts and complex optimized structures through TO [13], the authors’ focus in this paper is in determining the process constraints of the PolyJet process. Specifically, the authors first identify four key PolyJet manufacturing constraints that could affect part design via TO (Section 2.3). The results of experimentation to identify the geometric and process variables that have a statistically significant effect on each of the manufacturing considerations is then presented.

The results of experimentation to identify the geometric and process variables that have a statistically significant effect on each of the manufacturing considerations is then presented. The four sources of process constraints are (i) support material removal (Section 3), (ii) minimum resolvable feature size (Section 4), (iii) feature survivability during cleaning (Section 5), and (iv) minimum self-supporting angle (Section 6). Motivation, experimental methods, results, and discussion are included for each of the four manufacturing considerations. Conclusions and recommendations for future work are offered in Section 7.2. DFAM FOR POLYJET 3D PRINTING2.1. PolyJet OverviewOf particular interest in this study is multi-material jetting (e.g., Stratasys’ PolyJet technology), due to its ability to recreate fine geometric design features and its unique multi-material capabilities.

Material jetting PolyJet printing is an AM material jetting process, wherein droplets of liquid photopolymer are deposited directly onto an elevator via a series of inkjet printheads [14]. As the material is deposited, two ultraviolet (UV) lamps cure the photopolymer in multiple passes. Each subsequent layer is jetted on top of the previous one. A representation of this process can be seen in Figure 1.Figure 1. Representation of Direct 3D PolyJet Printing ProcessThe PolyJet process is able to create small features as it has a layer thickness of 16-32 microns and an in-plane resolution of 42 microns. In addition, the PolyJet process offers one significant, unique advantage among modern additive manufacturing process: the PolyJet process is capable of depositing two different materials on a pixel-by-pixel basis. This allows for creation of monolithic structures from multiple materials in a single build.

This allows for creation of monolithic structures from multiple materials in a single build. One material combination of engineering relevance is VeroWhite+, a rigid, white plastic-like material, and TangoBlack+, an elastomeric, flexible black material. The two materials can be combined in various ratios to create nine gradient material blends with elastic modulus properties ranging along the continuum of the two extremes. The Polyjet process also uses a sacrificial, hydrophobic support material during part creation. After printing, this gel-like photopolymer must be manually removed using the recommended water jet cleaning station.2.2. Prior Characterization of PolyJetNo researchers have yet created a detailed guide of DfAM considerations for the PolyJet process. Likewise, few have attempted to determine the effects that the PolyJet’s various parameters (printhead resolution, support material usage, multi-material options, etc.) have on these considerations.

Likewise, few have attempted to determine the effects that the PolyJet’s various parameters (printhead resolution, support material usage, multi-material options, etc.) have on these considerations. A review of existing work shows that i) existing studies tend to focus on material characterization, ii) process/property relationships have not been fully explored, and iii) the key process constraints of PolyJet have yet to be identified.Regarding material characterization, Moore and Williams investigated the fatigue life of the elastomeric TangoBlack+ material produced by the PolyJet process [15]. They were able to create an experimental curve relating the percent elongation of a TB+ specimen to the expected life of the specimen. They also derived several design rules from their study; one example is to avoid necks in a designed part, since necks were found to reduce the fatigue life of specimens.

They also derived several design rules from their study; one example is to avoid necks in a designed part, since necks were found to reduce the fatigue life of specimens. Pilipovic and co-authors measured the tensile and flexural properties of VeroBlack, VeroBlue, and Fullcure 720 materials; they found that the Fullcure 720 performed the best of the three, with a flexural strength slightly higher than the manufacturer’s quoted value [16]. The Fullcure material was also found to have a tensile strength lower than the manufacturer’s cited value. Vieira and co-authors found that the flexural modulus of Fullcure 720 can be further improved through UV or thermal post-processing, though this also has anegative effect on fatigue life [17]. Finally, Gibson and coauthors investigated the elastomeric properties of Fullcure 720 in both compression as well as bending [18].

Finally, Gibson and coauthors investigated the elastomeric properties of Fullcure 720 in both compression as well as bending [18]. Their experimentation showed that, while the material can be used for the creation of “living hinge” designs, its elastomeric properties are not strong enough to withstand heavy use.In addition to material characterization, there has also been some prior work into characterizing the PolyJet process and its effect on final parts. Udroiu investigated the surface finish of parts made of Fullcure 720 and found that the machine’s glossy surface finish setting gave parts a much lower surface roughness than the matte setting [19,20]. He also performed a basic investigation into how build time and material consumption changes relative to part orientation. From this investigation, he was able to provide several guidelines for optimal part orientation (e.g., “the smallest dimension along z-axis” and “highest part to the left”).

From this investigation, he was able to provide several guidelines for optimal part orientation (e.g., “the smallest dimension along z-axis” and “highest part to the left”). Singh likewise characterized the material consumption and build time of a nominal test part with the PolyJet process and confirmed Udroiu’s claim that the largest part dimensions should be aligned in the X-axis [21,22]. In addition, he investigated the dimensional accuracy of the part when built in different orientations and found that the x-axis orientation also offered the best dimensional accuracy for parts.Kim and Oh also offer an investigation of PolyJet material and system characterization. In their paper, they investigate numerous properties of manufactured PolyJet parts [23]. The authors discovered a 50% decrease in PolyJet tensile strength in the z-direction of parts compared to the x-direction.

The authors discovered a 50% decrease in PolyJet tensile strength in the z-direction of parts compared to the x-direction. They also discovered that the tensile strength of PolyJet parts decreases steadily and significantly as temperature is increased. While Kim and Oh did investigate minimum feature size, one of the goals of the research proposed in this paper, they did so only with simple rib and rectangular/circular gap shapes in a glossy finish. These shapes are capable of providing a benchmark minimum feature size, but may not capture the variation in size that may come from alternate orientations, surface finishes, and feature shapes (for example, the droplet-based nature of the PolyJet may allow it to recreate positive circular features more effectively than rectangular ones).Investigation of the properties at the interface of disparate printed materials is one of the potential key focuses of PolyJet research.

In addition to their study of general PolyJet material fatigue life, Moore and Williams also investigated the effect that the presence of a sharp material interface would have on the fatigue failure location of a test specimen [15]. They found that having a sharp interface between TangoBlack+ and VeroWhite+ loaded in shear would almost always cause the specimen to fail at the interface (92% of their specimens failed at the shear interface while only 8% failed at the weaker TB+ material). However, despite being a central tenet behind the creation of PolyJet parts, the multi-material interface is still generally under-investigated in existing literature.As this review demonstrates, PolyJet characterization research has included some investigation into manufacturing considerations such as material anisotropy and feature size. However, each of these studies tend to focus on one or two constraints and investigate the effects of only one variable on the constraints.

However, each of these studies tend to focus on one or two constraints and investigate the effects of only one variable on the constraints. A broader approach is needed to establish a larger field of PolyJet manufacturing constraints and to identify the essential process/geometry variables for each constraint. The experimental methods (Sections 3-6) of this work seek to fill this gap in understanding.2.3. Identification of PolyJet Manufacturing ConsiderationsTo illustrate the importance of manufacturing constraints when combined with topology optimization, consider the optimized compliant inverters from [13] (Figure 2). Both were designed using TO methods and subsequently printed via PolyJet AM.Figure 2. Optimized Compliant Inverters Manufactured via PolyJet AM [13]While the geometry of these inverters is relatively simple, they nevertheless help to illustrate certain concerns and possibilities that may arise in the use of automated design tools in AM.

These potential manufacturing constraints are identified as follows (note that the callouts in Figure 2 match the numbering below):1. Support Material Removal: Depending on part orientation and surface finish, the triangular holes in the single-material specimen would be filled with the PolyJet process’s hydrophobic support material. It is necessary to ensure that these holes are large enough to allow for rapid, manual removal of the support material during cleaning.2. Minimum Feature Size: In the case of the compliant inverter design domain, the TO algorithm tends to create narrow, hinge-like sections in the final design in order to increase deflection at the tip. Knowing the minimum manfacturable feature size in the PolyJet process, and incorporating that value into the TO formulation, will allow these joints to be a small as possible. This in turn will increase the maximum deflection of the inverter.3.

This in turn will increase the maximum deflection of the inverter.3. Survivable Feature Size: As mentioned, PolyJet support material must be removed manually with a high-pressure water jet. Experience shows that it is possible for smaller features (such as the long, thin arms and tip of the inverter) to break during the cleaning process. Because of this, it is crucial to determine the factors that drive part breakage and establish the minimum feature size that can survive the cleaning process.4. Self-Supporting Angle: To remove any concerns about support cleaning or feature survivability, it is possible to completely eliminate support material use in the creation of PolyJet parts. By doing this, it becomes necessary to ensure that angled faces (such as the one highlighted with the dashed line) will not collapse without support material.

By doing this, it becomes necessary to ensure that angled faces (such as the one highlighted with the dashed line) will not collapse without support material. This can be predicted by determining the self-supporting angle of angled faces.The experimentation that follows in the remainder of this paper will focus on identifying the geometric and process variables that have a statistically significant effect on these four manufacturing constraints. Without first understanding which parameters have an effect on the constraints, it is difficult (if not impossible) to establish a robust set of quantitative DfAM relationships to incorporate into TO.3. SUPPORT MATERIAL REMOVAL3.1. MotivationAs mentioned in Section 2.1, the PolyJet process uses a hydrophobic, gel-like photopolymer as a secondary sacrificial support material.

SUPPORT MATERIAL REMOVAL3.1. MotivationAs mentioned in Section 2.1, the PolyJet process uses a hydrophobic, gel-like photopolymer as a secondary sacrificial support material. As in other AM processes, this support material is used to ensure overhanging and angled structures are built properly (as was mentioned in reference to the triangular gaps in Figure 2). As the support material also affects a part’s surface finish, the PolyJet process offers two surface finish options during pre-processing: selecting the “matte finish” provides a uniform surface finish to the part by coating every surface in support material (even top-facing surfaces); selecting the“glossy finish” only provides support material to overhanging features (typically resulting in the top- facing surfaces having a more glossy finish than the remainder of the part). This support material is then removed manually with small tools or with a high-pressure water jet.

This support material is then removed manually with small tools or with a high-pressure water jet. The need to manually remove support material places significant limitations on the final geometry from an automated design process. Any fully enclosed cavities will be filled completely with support material that cannot be removed. In addition, certain long, narrow cavities may be difficult to clean.The experiments described in this section are designed to answer the question, What process parameters and geometric features affect the ability to successfully remove support material from printed parts?3.2. Experimental MethodsWhile no research has been done in support material removal specifically for the PolyJet, Vayre and coauthors began an experimental investigation into the cross-sectional size necessary to remove powder support material from channels in parts manufactured with electron beam melting [24,25].

They used compressed air to remove loose support powder from cylindrical channels of varying diameters and measured the depth of powder removed after a specified period of time. They found that the as the cross- sectional area of a channel was increased, the depth of powder removed from it also increased. They also noticed a qualitative difference between the channel geometry and how much powder was removed (cylindrical tubes were more easily cleaned than square ones). Of note is that the authors mention the need to avoid fully enclosed internal cavities due to the presence of support material, which is a consideration that would be important to account for in automated design methods.A similar experiment is performed here where the PolyJet support material is removed with a water- jet.

The support-filled channel specimens for this experiment rely on variation of three independent variables: (i) channel geometry, (ii) channel cross-sectional area, and (iii) channel connectivity (open at both ends or closed at one end). These geometry and cross-sectional area variables are included since they were found to have importance by Vayre and coauthors, while connectivity is a new inclusion based on our existing knowledge of the PolyJet cleaning process. Each variable, its two states, and its associated hypothesis are presented in Table 1. For this experiment, the length of the channel is held constant (at an arbitrary length of 44.2 mm) as is the total cleaning time (set at 2 minutes). An example of the test specimens is shown in Figure 3. Two iterations of each chosen variable combination were tested.Figure 3. Example of Support Cleaning Test SpecimensTable 1.

An example of the test specimens is shown in Figure 3. Two iterations of each chosen variable combination were tested.Figure 3. Example of Support Cleaning Test SpecimensTable 1. Experimental Variables in Support Material Removal Independent VariablesVariable Name Cross-Sectional GeometryCross-Sectional AreaConnectivityState 1Circle3.14 mm2ClosedState 2Square19.64 mm2OpenRationale/Hypothesis Support material may prove more difficult to remove from the sharp corners of a square cross-section. It may prove more difficult for the water stream to break away the support material in a smaller area.

It may prove more difficult for the water stream to break away the support material in a smaller area. If the channel is closed at one end, support material may be removed more slowly since it will only have one exit point.Dependent VariableVariable NameSupport Material RemovedUnits% (by mass)Rationale Measuring the percent mass of support material removed will help to determine the variables that slow the cleaning process (or impede it entirely).To facilitate careful and consistent cleaning practice, a test rig was designed to maintain a constant distance and angle from the support-filled channels using a constant water pressure of 1750 psi (Figure 4). The specimens were each cleaned for two minutes using this rig; the amount of support material removed was then calculated.Figure 4.

The specimens were each cleaned for two minutes using this rig; the amount of support material removed was then calculated.Figure 4. Cleaning Test RigTo determine the amount of support material that is removed in each test specimen, an initial piece with known support volume, model volume, and model density is weighed to calculate the density of support material. The total percent of support material removed is determined by comparing the mass of the experimental specimens before and after cleaning. This metric is used to represent the ease of removal of support material for a given geometry; the higher the percentage of support material removed at the end of the two minutes, then the easier the geometry is to clean.3.3. ResultsANOVA analysis was performed on an L8 Taguchi array with two iterations of specimen cleaning. The average amount of support material removed across all specimens was 25.4% (by mass, seen in Table 2).

The average amount of support material removed across all specimens was 25.4% (by mass, seen in Table 2). This value offers a general benchmark of how much support material is removed at this particular time/size scale, but does don’t offer much information regarding the roles of the independent variables. The results in Table 3 offer more detail about the role of each variable. The information in the table suggests that only the cross-sectional area of the cleaning specimens affects the amount of support material removed in a fixed amount of time; it is the only variable with a p-value < 0.05. Using the larger cross-sectional area results in a higher percentage of support material removal, as hypothesized, with an approximate increase of almost 12% over the smaller cross-section. However, it is worth noting the relatively low R2 value when compared to the other experiments in this paper (0.742, from Table 1).

However, it is worth noting the relatively low R2 value when compared to the other experiments in this paper (0.742, from Table 1). Figure 5 shows the desirability predictions based on the ANOVA analysis; the square cross-section, larger cross-sectional area, and closed channel configuration is predicted to offer the highest percentageof support material removed (though, again, only the cross-sectional area is deemed statistically significant).Table 2. Summary of Fit – Support CleaningRSquare Root Mean Square Error Mean of Response Observations (or Sum Wgts)0.742 5.301 25.409 8Table 3. Parameter Estimates – Support CleaningTerm Intercept Geometry[Circle] Cross-Sectional Area[1] Connectivity[Closed]Estimate Std Error25.409 -2.331 -5.903 0.2331.874 1.874 1.874 1.874t Ratio Prob>|t| <.001* 13.56 0.282 -1.24 0.035* -3.15 0.907 0.12Figure 5. Desirability Prediction – Support Cleaning3.4.

Desirability Prediction – Support Cleaning3.4. DiscussionThe result that suggests that increased cross-sectional area providing easier support material removal aligns with intuition, as well as the results found by Vayre and coauthors. However, it is surprising to see that connectivity may play no part in support material removal (according to the ANOVA analysis). It is possible that the connectivity would play a larger role if cleaning time was also accounted for as an independent variable; as more material is removed, the water stream will reach the other end of the channel, which may increase the observed effect of the connectivity.4. MINIMUM RESOLVABLE FEATURE SIZE4.1. MotivationEvery AM process has a limit on the smallest feature that it can create. This is most closely tied to the size of the deposition nozzle (or beam diameter, or particle size), the layer thickness, and the resolution of the x, y, and z direction motors in the printer frame.

This is most closely tied to the size of the deposition nozzle (or beam diameter, or particle size), the layer thickness, and the resolution of the x, y, and z direction motors in the printer frame. Stratasys specifies that the PolyJet process has a resolution of 42 micron in the x- and y-directions with a layer thickness between 16 and 32 microns. However, it is necessary to validate how these quoted resolution values influence the smallest feature size that can be printed (which in turn relates back to the thin hinge like sections in the inverters from Figure 2). While there has been some preliminary investigation into this with simple single- material rib shapes, as seen in [23], it is important to understand if the minimum manufacturable feature size differs between VeroWhite+ and TangoBlack+ material, surface finish types (denoted in the printeras “matte” vs.

“glossy”), part orientation, feature direction (embossed or debossed), and feature geometry.The experiments described in this section are designed to answer the question, What processparameters and geometric features affect the smallest part feature size that can be created by PolyJet?4.2. Experimental MethodsA standardized test specimen has been proposed by authors from NIST for use in characterizing the products of AM processes (Figure 6) [26]. Each feature of the specimen is meant to directly relate back to an aspect of the manufacturing process (e.g., resolution of the XY motors, layer thickness, etc.).This specimen includes both rectangular and circular features to quantify the minimum feature size of the AM machine. These minimum feature aspects will inform the design of our own minimum feature specimen, created solely to investigate the PolyJet process’s ability to create small embossed and debossed features.Figure 6.

AM Feature Specimen Proposed by NIST [26]As the high resolution of the PolyJet process is one of its primary advantages, there was concern over whether or not the NIST specimen would be small enough to identify the uniquely high resolution of the PolyJet AM process. As such, a specimen was designed which maintains the same geometry as the NIST specimen, but utilizes even smaller lower limits on the feature geometry being printed. In addition, it uses significantly less material than the NIST specimen, which facilitates efficient experimental repetition. The specimen design, seen in Figure 7, has been divided into quarters; one quadrant of the specimen is printed based on what type of feature is being tested.The independent variables for this experiment include (i) material type, (ii) surface finish, (iii) orientation, (iv) feature direction, and (v) feature shape. These variables, their two states, and the hypothesis associated with each are shown in Table 4.

These variables, their two states, and the hypothesis associated with each are shown in Table 4. The dependent variable is the minimum feature recreated by the PolyJet process. All specimens were cleaned using an Objet-recommended water-jet station with a pressure of 1750 psi. Two iterations of each tested variable combination were performed.Figure 7. Minimum Feature SpecimenTable 4. Experimental Variables in Minimum Resolvable Feature SizeIndependent VariablesVariable NameMaterial TypeSurface FinishOrientationFeature DirectionShapeState 1TB+GlossyXYDebossedCircleState 2VW+MatteXZEmbossedRectangleRationale/Hypothesis The compressive force of the roller may deform small features in the flexible TB+ during printing, hampering creation of smaller features. The extra support material required for the matte surface finish may better support smaller embossed features and maintain the geometry of smaller debossed features.

The extra support material required for the matte surface finish may better support smaller embossed features and maintain the geometry of smaller debossed features. The PolyJet’s Z-resolution of 32 microns is actually finer than the X-Y resolution of 42 microns, so smaller features may be possible in the XZ orientation. The compressive force of the roller may damage embossed features during printing (but cannot damage debossed features). Cleaning may also damage smaller embossed features. The droplet-based deposition pattern of the PolyJet process may have difficulty recreating the sharp corners of rectangular cross-sections at small feature sizes.Dependent VariableVariable Name Minimum Feature SizeUnitsmmRationale Measuring the minimum feature size will determine the smallest feature allowed in an automated design algorithm.4.3.

ResultsANOVA analysis was used on a L12 Taguchi array to determine which of the independent variables has the greatest impact on the minimum resolvable feature size. A feature was considered resolvable if it maintained the true nature of the cross-sectional shape (e.g., rectangles had 4 distinct linear sides) and was of the same height as the other specimens (though this is difficult to validate with the debossed features, since measurement tools will not fit inside the extrusions). Because the minimum feature size was determined by human interpretation (with the assistance of a digital microscope), it is possible that the final statistical data might vary from observer to observer. Statistically significant values are denoted with an asterisk.From this data, the mean resolvable feature size is approximately 0.73 mm (Table 5).

Statistically significant values are denoted with an asterisk.From this data, the mean resolvable feature size is approximately 0.73 mm (Table 5). As with the average value from the support removal experiments, this number simply offers an approximate benchmark accounting for changes to all independent variables. More importantly, the results of experimentation show two statistically significant factors that contribute to the minimum PolyJet feature size. Shown in Table 6, these are the print material (VW+ or TB+, with a p-value of 0.002) as well as the feature direction (embossed or debossed with a p-value of 0.003). Print material has the greatest effect on feature size (potentially causing a change of 0.330 mm from one material to the other) while the feature direction has the second greatest effect (with a magnitude of change of 0.288 mm between debossed and embossed features). Both of these values can be derived from the estimates in Table 6.

Desirability Prediction – Minimum Feature4.4. DiscussionMuch was learned from observing the printed minimum feature specimens underneath the digital microscope. VeroWhite+ material proves to be most beneficial because it creates much more clear and defined geometry in small features than TangoBlack+ (the spongy nature of TB+ has a tendency to blur the boundaries of small features). In addition, XZ-oriented specimens had features which appeared “squashed” when viewed under a microscope (e.g., circular cross-sections appeared as ovals). This is most likely an effect of the poorer XY resolution discussed earlier; the resolution causes the dimensions of the circular cross-section to be manufactured larger in the X-direction, giving the squashed appearance. As discussed earlier in Section 1.2, the force of the waterjet cleaning tends to break smaller parts.

As discussed earlier in Section 1.2, the force of the waterjet cleaning tends to break smaller parts. Even though the embossed features still performed better than the debossed features, this type of breakage can be seen very clearly in these minimum feature size specimens. Specimens with matte, embossed features are especially susceptible to breakage, since there is more support material that needs to be removed from around the delicate embossed features. However, with the exception of the glossy XY specimens, all specimens require cleaning after printing (for consistency, the glossy XY specimens were cleaned as well, even though they had no support material). This means that it is impossible to tell the “true” minimum manufacturable feature size, since it is always possible that features will break during cleaning.

This means that it is impossible to tell the “true” minimum manufacturable feature size, since it is always possible that features will break during cleaning. Because of this, the next section (Section 5) will present experiments where different features were placed under load by the waterjet to determine the variables that affect a feature’s ability to survive the cleaning process.5. SURVIVABLE FEATURE SIZE5.1. MotivationAn interesting additional consideration for the PolyJet process is that, while the system may be able to reproduce a feature of an incredibly small size (as shown in the minimum resolvable feature size specimens), the feature may be broken while using the water jet to clean the part. This has already been alluded to with the long, thin arms and displacement tip on the two printed inverters shown earlier in Figure 2.

This has already been alluded to with the long, thin arms and displacement tip on the two printed inverters shown earlier in Figure 2. To ensure that a part can survive the cleaning process, it may be necessary for the smallestprinted feature to be larger than what the machine’s smallest resolvable feature size actually is. This section will focus on determining the variables that have a dominant effect on the survivability of any printed partsThe experiments described in this section are designed to answer the question, What process parameters and geometric features affect the smallest part feature size that can be cleaned without breaking?5.2. Experimental MethodsThe same rig from the support cleaning tests is used to determine the factors which affect minimum survivable feature size. However, the geometry of the specimens, and thus the specimen holder, differs slightly.

However, the geometry of the specimens, and thus the specimen holder, differs slightly. For this experiment, the feature under testing is kept simple; initial focus will be on investigating a beam-like feature, since this compares well to the geometry of the minimum feature specimens and the printed compliant inverters. The independent variables under investigation include (i) feature cross- sectional area, (ii) feature length, (iii) build material, and (iv) connectivity (e.g., cantilevered at one end or fixed at both ends). These variables will allow us to gain a general understanding of which factor most drives breakage during cleaning. Note that in this experiment, the variable “length” simply refers to the distance between the two ends of the printed feature.Representative specimens are shown in Figure 9. Survivability is considered binary (either the feature broke after the allotted 1-minute cleaning time or it did not break).

Survivability is considered binary (either the feature broke after the allotted 1-minute cleaning time or it did not break). This experiment differs slightly from the previous ones; a full factorial experiment of these three variables was performed twice: once using VW+ for all of the tests and once using TB+ for all tests. The variables, their possible states, and the hypothesis for each are collected in Table 7.Figure 9. Example of Cleaning Survivability Test SpecimensTable 7.

The variables, their possible states, and the hypothesis for each are collected in Table 7.Figure 9. Example of Cleaning Survivability Test SpecimensTable 7. Experimental Variables for Survivable Feature SizeIndependent VariablesVariable NameState 1State 2Rationale/HypothesisConnectivityFixedCantileveredHaving the feature cantilevered at only one end would result in a higher reaction load at that point, increasing the shear stress and the likelihood of part breakage at the joint.Cross- Sectional Area0.75 mm1.50 mmA smaller cross-sectional area will result in larger stresses in the feature due to the force of cleaning, which would again increase the likelihood of breakage.Length17.00 mm6.86 mmHaving a longer specimen should increase the length of the moment arm at the fixation points due to the force of the waterjet.

This will increase the normal stresses at the joint and increase potential for breakage.Dependent VariableVariable NameSurvivabilityUnitsYes or NoRationale Measuring survivability (in simple yes/no terms) will help to determine the variables that determine if a particular element in a final design will be able to survive the cleaning process.5.3. ResultsThe survivability specimen experimentation was first performed for 8 combinations of the main factors for VW+ material (based on an L8 Taguchi array). The results from the ANOVA analysis of this experimentation are as follows. Note that a score of 2 denotes a survivable feature, while a score of 1 denotes a broken specimen. The mean response of 1.438 in Table 8 shows that approximately 44% of the tested VW+ specimens fractured under the waterjet loading (a value of 1.000 would mean that all specimens broke).

The mean response of 1.438 in Table 8 shows that approximately 44% of the tested VW+ specimens fractured under the waterjet loading (a value of 1.000 would mean that all specimens broke). However, as the p-values in Table 9 demonstrate, almost all of the broken specimens are due to the cross-sectional area (p-value of 0.002). The desirability predictions in Figure 10 reinforce this notion; the predictions recommend a large cross-section along with fixed connectivity and a long length to minimize VW+ specimen breakage (though, as already mentioned, the effects of connectivity and length are not statistically significant).Table 8. Summary of Fit – Survivable Feature Size RSquare Root Mean Square Error Mean of Response Observations (or Sum Wgts)0.927 0.177 1.438 8Table 9.

Summary of Fit – Survivable Feature Size RSquare Root Mean Square Error Mean of Response Observations (or Sum Wgts)0.927 0.177 1.438 8Table 9. Parameter Estimates – Survivable Feature SizeTerm Intercept Connectivity[Cantilevered] Cross-Sectional Area[Large] Length[Long]Estimate Std Error1.438 -0.063 0.438 0.0630.063 0.063 0.063 0.063t Ratio Prob>|t| <.001* 23.00 0.374 -1.00 0.002* 7.00 0.374 1.00Figure 10. Desirability Prediction – Survivable Feature SizeAfter performing the experiment with VW+ material, the same 8 combinations were printed using TB+ material. When tested, every cantilevered TB+ specimen survived cleaning while every specimen fixed at both ends broke, regardless of length or cross-sectional area. This shows that connectivity should be the only significant independent variable when considering survivability of TB+ material. This differs from the VW+ specimens, which break almost entirely due to the specimen’s cross-sectional area.5.4.

This differs from the VW+ specimens, which break almost entirely due to the specimen’s cross-sectional area.5.4. DiscussionThe differing results between VW+ and TB+ specimens makes intuitive sense. For the more rigid VW+ material, thin features will not flex under the load of the water jet. Because of this, the cross- sectional area dominates whether a feature will survive, with larger areas experiencing lower shear stress from the waterjet load. This gives them a higher chance for survival (in the case of this experimentation,a diameter of 1.5 mm was consistently able to survive cleaning). On the other hand, when cantilevered, the elastomeric nature of the TB+ material causes it to flex under the water jet instead of fracturing. When the load is removed, the flexible material rebounds to its original position. When the TB+ specimen is fixed at both ends, it is unable to freely deflect and instead ruptures at the point of loading.6. SELF-SUPPORTING ANGLE6.1.

When the TB+ specimen is fixed at both ends, it is unable to freely deflect and instead ruptures at the point of loading.6. SELF-SUPPORTING ANGLE6.1. MotivationWhile the PolyJet process in its default configuration relies on the use of support material for any constructed angle (save for perfectly vertical faces), it is possible to turn off the use of support material in the system completely. This would allow a designer to create a specimen, such as the inverters from Figure 2, without worrying about 1) whether or not the triangular holes could be cleaned easily or 2) whether or not the waterjet would cause the fragile arms of the inverter to break. To take full advantage of this feature, it is important to know how certain process variables affect the self-supporting angle when support material has been eliminated.

To take full advantage of this feature, it is important to know how certain process variables affect the self-supporting angle when support material has been eliminated. By knowing understanding these effects, a designer could be confident that angled surfaces from a TO algorithm (such as the surface highlighted with a dashed line in Figure 2) could be created properly without compromising the geometry of the final part or the safety of the machine.The experiments described in this section are designed to answer the questions, What processparameters affect the most acute angle that can be created in the absence of support material?6.2. Experimental MethodsA simple specimen consisting of 10 faces at steadily increasing angles was designed to quantify the self-supporting angle of PolyJet parts when support material is turned off.

Experimental MethodsA simple specimen consisting of 10 faces at steadily increasing angles was designed to quantify the self-supporting angle of PolyJet parts when support material is turned off. The two independent variables under investigation are (i) material choice and (ii) orientation of the angled face with respect to the printhead’s roller. Orientation is tested to determine if the roller is potentially pushing deposited droplets off of the self-supporting edge. If the roller is pushing droplets, this could result in a poorer self- supporting angle when the motion of the roller is perpendicular to the front edge of the specimen instead of parallel. The specimen used for testing (seen in Figure 11) contains 10 faces, starting from a vertical 90 degree face (measured from the horizontal) and moving to a 72 degree face in 2 degree increments.

The specimen used for testing (seen in Figure 11) contains 10 faces, starting from a vertical 90 degree face (measured from the horizontal) and moving to a 72 degree face in 2 degree increments. The candidate orientations are 1) with the specimen’s longest dimension aligned in the x-direction and 2) with the specimen’s longest dimension aligned in the y-direction. These variables are collected in Table 10 along with their associated hypotheses. Each possible combination of variables was tested twice.Table 10. Experimental Variables in Self-Supporting AngleIndependent VariablesVariable NameState 1State 2Rationale/HypothesisMaterial TypeOrientationTB+X- DominatedVW+Y- DominatedThe compressive force of the roller may deform the flexible TB+ material during printing, causing fresh droplets to fall off the upper surface. If the material droplets have different viscosity properties, then one may be able to support higher angles than the other.

If the material droplets have different viscosity properties, then one may be able to support higher angles than the other. The roller may be pushing droplets off of the edge, which could result in a poorer self-supporting angle for a specimen predominantly oriented in the y-direction.Dependent VariableVariable NameSelf-Supporting AngleUnits Degrees (from Horizontal)Rationale Measuring the self-supporting angle with the given independent variables will show the recommended angle, orientation, and material type for printing optimized parts without support material.Figure 11. Self-Supporting Angle Specimen6.3. ResultsExperimentation shows that all specimens failed before the final 72 degree angled face. This failure is shown through stalagmites that appear to grow right beneath the angled face (an example is shown in Figure 12).Figure 12.

This failure is shown through stalagmites that appear to grow right beneath the angled face (an example is shown in Figure 12).Figure 12. Stalagmites Denote a Non-Self-Supporting Angled FaceThe maximum self-supporting angle was determined to be at the face before significant stalagmite growth was shown. The maximum angle was recorded for each of the 8 printed specimens. ANOVA analysis was performed on a full-factorial array to determine which of the factors, if any, had a significant effect on the maximum self-supporting angle.The mean benchmark self-supporting angle for all specimens is 84.5 degrees from horizontal (see Table 11). Again, this average value accounts for changes in both independent variables. From this analysis, it appears that material choice (p = 0.205) does not have a statistically significant effect (p < 0.05) on the maximum self-supporting angle (shown in Table 12).

From this analysis, it appears that material choice (p = 0.205) does not have a statistically significant effect (p < 0.05) on the maximum self-supporting angle (shown in Table 12). However, there does seem to be a potential pattern related to orientation that may merit further investigation with more test iterations (p = 0.064); the mean angle for the x-dominated specimens was 87 degrees from horizontal while the mean angle for y-dominated specimens was 82 degrees. While the p-value in Table 12 does not confirm statistical significance, more specimen iterations may decrease the standard deviation in the orientation data and better show if orientation has a significant effect. Based on the desirability predictions from this experiment (seen in Figure 13), the best self-supporting angle would be achieved on a specimen with TB+ material oriented in the y-direction.Table 11.

Summary of Fit – Self-Supporting Angle RSquare Root Mean Square Error Mean of Response Observations (or Sum Wgts)0.991 0.5 84.5 4Table 12. Parameter Estimates – Self-Supporting AngleTerm Intercept Material[TB+] Orientation[X-Dominated]Estimate Std Error84.5 -0.75 2.50.25 0.25 0.25t Ratio Prob>|t| 0.002* 338.00 0.205 -3.00 0.064 10.00Figure 13. Desirability Prediction – Self-Supporting Angle6.4. DiscussionIt is worth noting that the desirability predictions in Figure 13 are directly contrary to the initial hypothesis that specimens oriented in the x-dominated direction would see a better self-supporting angle. This data suggests that having the roller passing over the material droplets in the direction of the angled face’s leading edge is actually beneficial.

This data suggests that having the roller passing over the material droplets in the direction of the angled face’s leading edge is actually beneficial. Unfortunately, the authors cannot offer a concrete hypothesis at this time as to why this may be the case (though it is possible that it could be connected either to the presence of the roller or to the nature of the material jetting from the nozzles). Also, while the predictions recommend the TB+ material over the VW+, there is no statistically significant evidence from this testing to confirm that it would truly offer a better self-supporting angle.7. CONCLUSIONS AND RECOMMENDATIONS FOR FUTURE WORKThis work presents a series of designed experiments to determine the key parameters that influence four specific manufacturing considerations in the PolyJet process.

The variables that affect minimum feature size, support material removal, feature survivability, and self-supporting angles are identified as follows:Support material removal experimentation demonstrated a reliance solely on the cross- sectional area of the channel.The minimum resolvable feature size has been shown to rely on material choice and feature direction.Feature survivability during cleaning was tied to cross-sectional area (for VW+ features) and feature connectivity (for TB+ features).Finally, the self-supporting angle in the absence of support material was driven by the orientation of the surface with respect to the roller, with y-dominated specimens offering better self-supporting angles.Identifying these key variables is a crucial to incorporating PolyJet manufacturing considerations into TO methods.

Future work will focus on 1) establishing clear quantitative relationships between the key factors identified in this paper and their associated dependent variables, 2) determining the lower limits ofthe dependent variables due to the effects of the key factors, and 3) investigating manufacturing considerations that arise specifically when printing stiff and flexible materials in a monolithic structure. This will include investigation of the material properties in the interface regions between disparate material elements as well as a quantifying the effects of changing feature size on digital material properties.8. REFERENCES[1]Brackett, D., Ashcroft, I., and Hague, R., 2011, “Topology Optimization for Additive Manufacturing,” 22nd Annual International Solid Freeform Fabrication Symposium, pp. 348–362.[2] Watts, D. M., and Hague, R., 2006, “Exploiting the Design Freedom of RM,” 17th Annual[3][4][5][6][7][8]International Solid Freeform Fabrication Symposium, pp.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDesign for additive manufacturing: Simplification of product architecture by part consolidation for the lifecycleSamyeon Kim1*, Yunlong Tang1, David W. Rosen1,21 Digital Manufacturing and Design Centre, Singapore University of Technology and Design, 8 Somapah Road, Singapore 4873722 The George W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332 *Corresponding author. Email: samyeon_kim@sutd.edu.sgAbstractAdditive manufacturing (AM) can support the fabrication of the complex design and generate new design opportunities for improving products. To identify and leverage these opportunities, design studies in early product design stages are required.

To identify and leverage these opportunities, design studies in early product design stages are required. Since part consolidation is one of AM design potentials in conceptual and embodiment design stages, this study proposes a design method to reconceptualize existing product design in the context of part consolidation. Function requirements and physical relations between existing parts are used to investigate AM design potential and identify candidates for consolidation. After identification of consolidation candidates, function sharing between parts and modules is checked because they have high possibilities to be consolidated if they share the same functions. Furthermore, AM design potential is identified to help designers add value in part design. In order to support designers, it is required to link AM design potential to the part candidates in order to explore AM design benefits. A case study with motorcycles is performed to demonstrate the proposed method.

A case study with motorcycles is performed to demonstrate the proposed method. The AM design potential for the case study contains the lifecycle considerations related to fuel savings due to lightweight, and simplified and less expensive assembly operations due to simplified product architecture by part consolidation.IntroductionThe evolution of AM has been made by AM materials and technologies. The growth of AM is driven by market needs, such as increasing demand of a customized product, manufacturing time and cost reductions, and shorter product development process [1]. The objective of Design for Additive Manufacturing (DFAM) has focused on both maximizing product performance [2] and alleviating manufacturing constraints of AM to improve manufacturability [1] in order to take advantages of AM. In order to explore AM benefits, DFAM research has been studied in various design stages.

In order to explore AM benefits, DFAM research has been studied in various design stages. In conceptual and embodiment design stages, previous studies provided a framework to guide designers to generate new design for AM and identify AM potential [3]. Function structure was determined based on product requirements and then both economic and technical feasibilities were considered to evaluate new design [3]. In detailed design, design optimization has been mainly concerned to improve performance, such as topology optimization and application of cellular structure [4, 5]. Although the optimized part design has high shape complexity, AM can fabricate the optimized part design. Accordingly, DFAM in detailed design has been researched to add values in the part design and provide motivations of adopting AM in industry. However, AM processes also have manufacturing constraints.

However, AM processes also have manufacturing constraints. Accordingly, benchmark studies have been performed to alleviate AM process constraints by providing design rules [6, 7]. Furthermore, DFAM has been focused on optimizing build orientation to control support structure generation and redesigning parts to avoid support structure generation [8].In these product design stages, design considerations, such as product, sustainability, and business considerations, should be considered as early as possible to leverage AM design potential. The product consideration is to improve3assembly operations by applying DFAM so that part consolidation and multi-part mechanisms without further secondary assembly operations are main considerations [9]. The sustainability consideration is to reduce harmful environmental impact by analyzing lifecycle of AM printed parts and then redesigning the parts for enhancing reusability and recyclability.

The sustainability consideration is to reduce harmful environmental impact by analyzing lifecycle of AM printed parts and then redesigning the parts for enhancing reusability and recyclability. Business consideration is to understand various costs from AM processes before adopting AM.Therefore, the scope of this study is for conceptual and embodiment design stages, which are less considered in previous studies. Since part consolidation (PC) is one of the best AM design benefits by integrating multiple parts and assembly interfaces including moving mechanisms, this study provides a design method to reconceptualize existing product designs in the context of PC. The reconceptualization aims to add value in product design and identify design opportunities for PC when adopting AM. The proposed method consists of three steps. Firstly, function diagram is identified to analyze functional and physical structures of existing product.

The proposed method consists of three steps. Firstly, function diagram is identified to analyze functional and physical structures of existing product. Secondly, important functions and function flows in the function diagram are selected to prioritize functions and flows. Lastly, parts with high important function and selected functional flows are selected and then grouped as candidates for consolidation. For generating new design concepts, physical relations among parts and function sharing between parts and modules are considered to group parts into larger modules for consolidation. The parts in the modules can be candidates for PC. Furthermore, AM design potential has been identified to support designers to leverage AM benefits while reconceptualizing the selected parts and modules.

Furthermore, AM design potential has been identified to support designers to leverage AM benefits while reconceptualizing the selected parts and modules. A case study of reconceptualizing existing motorcycle architectures as a novel electric motorcycle that are fully printed by AM is performed to demonstrate the usefulness of the proposed method.Literature reviewPC is a promising design approach to take advantages of AM in conceptual and embodiment design stages. Once parts for the consolidation are decided, design optimization methods for the parts can be applied to achieve multiple design objectives. Accordingly, previous studies of part consolidation have focused on part identification for PC. Another reason why the part identification receives attention is the limitation of the build chamber size of AM machines.

Another reason why the part identification receives attention is the limitation of the build chamber size of AM machines. AM cannot fabricate an entire product consisting of various parts with different sizes and materials but fabricate a part or module that have smaller than the size of the build chamber. Accordingly, identification of part candidates for PC is necessary. Accordingly, Yang, et al. [10], Yang and Zhao [11], Yang, et al. [12] developed rule-based design methods for design feasibility to identify the part candidates. The rules aim to reflect AM machine limitations, such as material availability and part size limitation, assembly considerations by checking relative motion between parts, and maintenance frequency. It would result in feasible sets of parts for PC. However, the rule-based methods have less considered how to utilize AM design potential and (re)design existing design as new design when adopting AM. In order to consider AM design potential, Lindemann, et al.

In order to consider AM design potential, Lindemann, et al. [13] provided a design methodology to consider the design feasibility and identify AM design potential, such as part complexity and part consolidation possibility. The design methodology helps assess the part candidates based on criteria for design feasibility and AM design potential and then select parts with the biggest design benefits for AM. Furthermore, it provides a process of redesigning selected parts based on function diagram but does not clearly describe the redesign process for generating new design. On the other hands, Kumke, et al. [14] proposed new opportunistic DFAM method which aims to support designers to leverage AM design potential. The DFAM method provided classification of AM design potential to understand interrelations between design potentials by using terms levers and value propositions.

The DFAM method provided classification of AM design potential to understand interrelations between design potentials by using terms levers and value propositions. The levers are several types of design freedom given by AM, while the value propositions are detailed benefits of the levers for customer as shown in Figure 1.

The interdependencies between the levers and the value propositions in the classification help identify promising design ideas based on existing product and new product design requirements.4AM design potentialLeversValue propositionsNon-uniform wall thicknessenablesLightweight designIntegrated designReliability improvementUndercuts are availableAssembly cost reductionFigure 1 Classification of AM design potential (reproduced from [14])From previous studies about PC, although many studies have dealt with functional analysis to understand functional and physical relationships between parts in order to decide design boundaries for PC, main contributions of them are identification of part candidates for consolidation rather than redesigning existing design to new design or generating new design by taking advantages of AM.

Few studies have introduced finding AM enabled solutions which is related to redesign and new design in the conceptual and embodiment design stages. It would be critical for DFAM to guide designers to generate new design concepts or reconceptualize existing design as new design by incorporating the designers and various AM design potential. Therefore, this study proposed a design method to focus on redesign and new design in the early design stage.Part consolidation method by reconceptualizationIn this study, a design method is proposed to support part consolidation in the conceptual and embodiment design stages. The objective of the proposed design method is to guide designers to reconceptualize existing product designs that only AM can achieve in the context of part consolidations. The proposed design method helps integrate selected parts, modules, and functions that are candidates for PC due to AM unique capabilities.

The proposed design method helps integrate selected parts, modules, and functions that are candidates for PC due to AM unique capabilities. This study provides three steps as shown in Figure 2.~ Existing p:oductAnalyze an existing pmductStep L Define functional/Physical relationsPart consolidationStep 2_ Decide important flowDesign optimizationStep 3_ Identify candidates for part consolidationFigure 2 A process of part consolidation method5Step 1. Define functional/physical relations Firstly, the functional diagram of existing product is identified as fundamental information in this study. The function diagram has function elements, functional flows, and physical chunk. Main functions of the existing product can be decomposed into functional elements in Figure 3. The functional elements will be used to identify parts that share same function as candidates for PC in this study.

The functional elements will be used to identify parts that share same function as candidates for PC in this study. Functional flows consist of material, signal, and energy flows between the functional elements so that they represent function types of the relations. It is important to group function elements to figure out physical chuck. Physical chuck is the design solutions to the functional elements. In other words, the functional elements can be matched with possible solutions which are concepts of parts, modules or systems. For example, as shown in Figure 4, function ‘store energy’ can be divided into specific functional elements and then matched with possible solutions like battery or fuel cell for ‘store electrical energy.’Human EnagyGasolineHuman Enagy.. .. Fudtank .....................................................__ , I Import H Store' OCtirolllf valve -: Regulate liquidEngineCovert liquid to h.e.1➔ SupplyHI-liquidliquid liquid .,.

Decide important flows Secondly, important flows are selected from the functional flows. It aims to identify functions or physical chunks that are closely linked by the same functional flows. When there are many parts in a product with multiple functions, the function diagram would be complex. Function elements can be connected by various functional flows between other elements. Accordingly, identifying the most important functions and flows is helpful to screen the elements to select candidates for PC. When selecting different important flow, the candidates for PC will differ. For example, when a material flow is selected as an important flow in Figure 3 for changing energy source of motorcycle from gasoline to electricity, five functions from ‘import liquid’ to ‘convert liquid to heat energy (h.e.)’ are selected.

When selecting energy flow starting from human energy, three functions ‘rotate steering’, ‘transfer rotational energy (r.e.)’, and ‘steer vehicle’ are selected as candidates for PC.6Step 3. Identify candidates for part consolidation Lastly, functions and parts that are linked to the importance flows are grouped into modules for PC. The functions and parts in modules have high possibility to be integrated, which is groups of candidates for PC. When integrating them, especially the parts, function sharing between the parts are considered to facilitate decision making for designers. If parts on the same important flow have common functions, the common functions can be shared by consolidating the parts.After that, design values that are inherited by AM should be added into the candidates. Adding design values is a starting point of reconceptualization by exploring AM benefits.

Adding design values is a starting point of reconceptualization by exploring AM benefits. However, due to the inexperience of AM, designers would face limitation of creating novel design and facilitating AM unique capabilities. Accordingly, these design values are necessarily identified to support designers to add design values in three different levels as shown in Figure 5. Design values in product level can contain mass customization, performance improvement, and production improvement. The values in the product level can be more specified into manufacturing level and part level. Design values in manufacturing level can deal with production efficiency and supply chain benefits, while values in part level focus on design benefits from high design freedom by AM, such as lightweight structure.

If more specific design values are listed and visualized, designers can utilize more diverse AM capability for design.MFG levelReduce assembly time and costProduct levelPart levelReduce part countsLightweightTopology optimizationValues added by AMImprove performanceImprove performanceCellular structureConformal cooling channelCustomized productImprove reliabilityFigure 5 Design value structure that AM can provideCase studyA case study is performed to demonstrate the proposed design method on conventional motorcycle architectures by reconceptualizing the architecture to take advantage of novel function sharing and part consolidation opportunities in the conceptual design of AM fabricated electric motorcycles. New concept design of the electric motorcycle in Figure 6 (b), called as NERA, is printed by BigRep that is a large scale AM company [16]. The purpose of fabricating the NERA is to show AM fabrication capabilities.

The purpose of fabricating the NERA is to show AM fabrication capabilities. When designing the NERA, designers arranged core parts that are relevant functionally and then decided to leverage AM design potential and avoid manufacturing constraints [16]. The design process of the NERA has similarity with the proposed design method. Accordingly, this case study is performed to show how existing design of the motorcycle in Figure 6 (a) can be reconceptualized as new concept design of an electric motorcycle that will be fully fabricated by AM in Figure 6 (b).7(a)(b)Figure 6 Reconceptualizing from (a) existing motorcycle (Courtesy of Ducati) to (b) electric motorcycle that is fully printed by AM (Courtesy of BigRep)Step 1 Define functional/physical relationships Main functions of the existing motorcycle are power generation for movement, steering direction, and vibration absorbing for comfort.

Original complex function diagram can be simplified by only considering these main functions as shown in Figure 3. The function diagram can be represented more specifically by using part-function diagram in Figure 7. In the part-function diagram, main modules and parts are determined and the modules are connected by functional flows and physical relations.Figure 7 Part-function diagram of the existing motorcycleStep 2 Select important flow and Step 3 Identify candidates for part consolidation Physical relations are selected as important flow to achieve part consolidation. As an example, parts related to steering the motorcycle, such as a wheel, tire, suspension, spoke, and steering handle, are selected based on their physical relations as shown in Figure 8. Additionally, functions of parts, such as ‘rotate wheel’ and ‘steering the motorcycle,’ are defined to check function sharing between parts.

Additionally, functions of parts, such as ‘rotate wheel’ and ‘steering the motorcycle,’ are defined to check function sharing between parts. If functions are shared between parts, such as front wheel and tire with the same function ‘rotate wheel,’ then these parts can be worth integrating each other. By considering physical relations and function sharing, in this case, two groups are determined: the group with the wheel and tire, and the group with the suspension, spoke, and steering handle. It means that parts in each group have high chance to be consolidated by joining assembly interfaces.8After selecting part candidates and before reconceptualizing existing part candidates, designers should understand AM benefits in Figure 5 to add values in the existing design for reconceptualization. In this case, new design of the electric motorcycle, the NERA, takes advantages of mainly part count reduction and lightweight values.

In this case, new design of the electric motorcycle, the NERA, takes advantages of mainly part count reduction and lightweight values. Accordingly, existing tire and wheel in the first group are consolidated and then reconceptualized by applying cellular structure as shown in Figure 8. It results in new concepts, which is airless tire and wheel. Parts in the second group from existing design are also reconceptualized by consolidating three parts into the forkless steering as shown in Figure 8.Physical relation Functional relationFunctionRotate wheelAbsorb vibrationSteering the motorcyclePartsWheel FTire FSuspension FSpokeSteering handleTire + Wheel (front)Suspension + Spoke + HandleFigure 8 An example of part consolidation based on physical relations and function sharingFor integrating functions, energy flow has chosen as important flow in Figure 3.

We focused on two functions of the existing motorcycle: ‘rotate wheel’ and ‘absorb vibration’ to add values for AM in Figure 9 (a). Then, parts that are related to these functions are determined such as rear wheel and tire, suspension, and chassis as shown in Figure 9 (a). Furthermore, it is checked whether any functions in these parts are shared. If the functions are shared, there is high possibility to integrate function. Therefore, we checked specific functions of suspension and chassis in Figure 9 (b). The chassis has multiple functions including ‘absorb vibration’, while the suspension has the only function ‘absorb vibration’. In this case, the function ‘absorb vibration’ is shared with two parts so that they can be redesigned by adding design values, such as cellular structure for lightweight structure, while consolidating each other. Finally, the proposed method shows possibility to help come up with novel body design of NERA by integrating same function from different parts.

Finally, the proposed method shows possibility to help come up with novel body design of NERA by integrating same function from different parts. The novel body design has the bumper with cellular structure, instead of using the suspension, for absorbing vibration and making the motorcycle light.9SuspensionRotate wheelAbsorb vibration(a)(b)Figure 9 Novel body design from function integration: (a) function sharing between parts, (b) function integrationTo sum up this case study, parts of conventional product are grouped into five modules in Figure 10 based on function sharing and selection of important flows. Since design team of the NERA aims to take advantage of part consolidation and lightweight structure, designers redesign the parts in the modules as new concepts for the NERA by considering these specific design benefits, such as ‘reduce assembly time and cost’, ‘improve reliability’, ‘lightweight’, and ‘improve performance’ as shown in Figure 5.

As shown in Figure 10, modules 1 and 2 are redesigned to airless tire and wheel and forkless steering by adding design values such as cellular structure for lightweight as described in Figure 8. For the same reason, modules 3 and 4 are redesigned as motor-embedded wheel. Since module 4 is related to engine, it is replaced to the motor and then consolidated with module 3. Module 5 related to engine should be replaced to battery and motor for the NERA. Novel body design in Figure 9 and Figure 10 is derived from function sharing between suspension and chassis. Since the novel body design contains cellular structure that has function of absorbing vibration, the original suspension is not required for the NERA.PC, cellular strnctmeMotor-embedded wheelModule rlacement10Batte1y MotorOverall, adding the design values for adopting AM in conceptual and embodiment design stages is critical because it can influence the lifecycle of a product.

By adding design value for the lightweight and consolidating parts for part count reduction, the reconceptualized design would reflect the lifecycle consideration potentially. Due to lightweight structure, less material is consumed in manufacturing phase and fuel can be saved in use phase of the lifecycle of the NERA. The part count reduction by PC has an impact on simplified and less expensive assembly operations in manufacturing phase and facilitating disassembly operation because of less disassembly joints in the end-of-life phase.Concluding remarksPart consolidation is one of the best ways to explore and take advantages of AM. This study provides a design method to support part consolidation when adopting AM. Especially, in order to explore AM capabilities in product design, the proposed method focused on the reconceptualization of existing design towards new design that only AM can fabricate.

Especially, in order to explore AM capabilities in product design, the proposed method focused on the reconceptualization of existing design towards new design that only AM can fabricate. For the reconceptualization, physical and functional relations are utilized to group parts into modules. Parts in the modules are considered as candidates for part consolidation. After selecting parts for the consolidation, the possible design values from AM are identified to help designers explore AM benefits during part consolidation. This study contributes to guiding designers to create new design by identifying candidates for part consolidation and reconceptualizing existing design in conceptual and embodiment design stages. Future work will be to model the function diagram in ontology and develop semantic network for AM design potential. It will allow easy identification of both function sharing between parts in the function diagram and AM design potential to generate promising AM design.

It will allow easy identification of both function sharing between parts in the function diagram and AM design potential to generate promising AM design. Since the AM design potential such as part consolidation and lightweight structure has an impact on the lifecycle, the lifecycle considerations will be more studied and implemented in the conceptual and embodiment design stages.Acknowledgement The authors acknowledge support from the Digital Manufacturing and Design (DManD) research center at the Singapore University of Technology and Design, projects funded by the National Additive Manufacturing Innovation Cluster (NAMIC) in Singapore, and Moveon Technologies Pte Ltd. Both DManD and NAMIC are supported by the Singapore National Research Foundation.Reference[1] Thompson, M. K., Moroni, G., Vaneker, T., Fadel, G., Campbell, R.

Both DManD and NAMIC are supported by the Singapore National Research Foundation.Reference[1] Thompson, M. K., Moroni, G., Vaneker, T., Fadel, G., Campbell, R. I., Gibson, I., Bernard, A., Schulz, J., Graf, P., Ahuja, B., and Martina, F., 2016, "Design for Additive Manufacturing: Trends, opportunities, considerations, and constraints," CIRP Annals - Manufacturing Technology, 65(2), pp. 737-760.[2] Rosen, D. W., 2014, "Research supporting principles for design for additive manufacturing," Virtual andPhysical Prototyping, 9(4), pp. 225-232.[3] Kumke, M., Watschke, H., and Vietor, T., 2016, "A new methodological framework for design for additivemanufacturing," Virtual and Physical Prototyping, 11(1), pp. 3-19.[4] Rosen, D. W., 2007, "Computer-Aided Design for Additive Manufacturing of Cellular Structures," Computer-Aided Design and Applications, 4(5), pp. 585-594.[5] Gibson, I., Rosen, D.

3-19.[4] Rosen, D. W., 2007, "Computer-Aided Design for Additive Manufacturing of Cellular Structures," Computer-Aided Design and Applications, 4(5), pp. 585-594.[5] Gibson, I., Rosen, D. W., and Stucker, B., 2015, Additive Manufacturing Technologies: 3D Printing, RapidPrototyping, and Direct Digital Manufacturing, Springer US, New York, USA.[6] Kruth, J. P., Vandenbroucke, B., Van Vaerenbergh, J., and Mercelis, P., 2005, "Benchmarking of different SLS/SLM processes as rapid manufacturing techniques," International Conference Polymers & Moulds Innovations (PMI)Ghent, Belgium, p. 525.[7] Rebaioli, L., and Fassi, I., 2017, "A review on benchmark artifacts for evaluating the geometrical performance of additive manufacturing processes," The International Journal of Advanced Manufacturing Technology, 93(5), pp. 2571-2598.[8] Zhang, Y., Bernard, A., Harik, R., and Karunakaran, K.

2571-2598.[8] Zhang, Y., Bernard, A., Harik, R., and Karunakaran, K. P., 2017, "Build orientation optimization for multi- part production in additive manufacturing," Journal of Intelligent Manufacturing, 28(6), pp. 1393-1407.11[9] ISO / ASTM52910, 2017, "Standard Guidelines for Design for Additive Manufacturing," ASTM International,West Conshohocken, PA.[10] Yang, S., Santoro, F., and Zhao, Y. F., 2018, "Towards a Numerical Approach of Finding Candidates forAdditive Manufacturing-Enabled Part Consolidation," Journal of Mechanical Design, 140(4).[11] Yang, S., and Zhao, Y. F., 2018, "Additive Manufacturing-Enabled Part Count Reduction: A LifecyclePerspective," Journal of Mechanical Design, 140(3).[12] Yang, S., Santoro, F., Sulthan, M. A., and Zhao, Y. F., 2019, "A numerical-based part consolidation candidate detection approach with modularization considerations," Research in Engineering Design, 30(1), pp. 63-83.

A., and Zhao, Y. F., 2019, "A numerical-based part consolidation candidate detection approach with modularization considerations," Research in Engineering Design, 30(1), pp. 63-83. [13] Lindemann, C., Reiher, T., Jahnke, U., and Koch, R., 2015, "Towards a sustainable and economic selectionof part candidates for additive manufacturing," Rapid Prototyping Journal, 21(2), pp. 216-227.[14] Kumke, M., Watschke, H., Hartogh, P., Bavendiek, A.-K., and Vietor, T., 2018, "Methods and tools for identifying and leveraging additive manufacturing design potentials," International Journal on Interactive Design and Manufacturing (IJIDeM), 12(2), pp. 481-493.[15] Eppinger, S. D., and Ulrich, K. T., 2003, Product design and development, New York: McGraw-Hill. [16] BigRep GmbH, 2019, "NERA e-motorbike," from https://bigrep.com/posts/deeper-look_into-the-fully-3d-printed-e-bike-nera/, Accessed 25 July, 2019.12.

Yang1, O. L. A. Harrysson2, D. Cormier3, H. West2, S. Zhang1, H. Gong4, B. Stucker5 1Department of Industrial Engineering, University of Louisville, Louisville, KY 40292 2Department of Industrial & Systems Engineering, North Carolina State University, Raleigh, NC 27695 3Department of Industrial & Systems Engineering, Rochester Institute of Technology, Rochester, NY 14623 4Department of Manufacturing Engineering, Georgia Southern University, Statesboro, GA 30458 53DSIM, 1794 Olympic Pkwy, Park City, UT 84098Abstract The design of lightweight structures realized via additive manufacturing has been drawing considerable amount of attentions in academia and industries for a wide range of applications. However, various challenges remain for AM lightweight structures to be reliably used for these applications.

However, various challenges remain for AM lightweight structures to be reliably used for these applications. For example, despite extensive advancement with geometric design, there still lacks adequate understanding with the process-material property relationship of AM lightweight structures. In addition, a more integrated design approach must also be adopted in order to take non-uniform material design into consideration. In our works, a design approach based on unit cell cellular structure was taken in the attempt to establish a comprehensive design methodology for lightweight structures. Analytical cellular models were established to provide computationally efficient property estimation, and various design factors such as size effect, stress concentration and joint angle effect were also investigated in order to provide additional design guidelines.

In addition, it was also found that the geometry and microstructure of the cellular structures are dependent on both the process setup and the feature dimensions, which strongly support the argument to establish a multi-scale hierarchical cellular design tool.Keywords: lightweight structure, additive manufacturing, cellular structure, unit cell, designIntroduction One of the objectives of structural design is to minimize the mass consumption and maximize the utilization efficiency of the materials. Therefore, lightweight structure design has always been sought after for almost all the engineering designs. Lightweighting brings about various technical advantages such as high strength to weight ratio, high energy absorption per weight ratio, low thermal conductivity, and large surface area to volume/weight ratio.

These attributes could in turn translate into various economical and environmental benefits such as product reliability, system energy efficiency and product sustainability. However, as lightweight designs often involve high level of geometrical complexity, the realization of these designs has been a challenging task with traditional manufacturing technologies. It has been widely recognized that additive manufacturing (AM) technologies possess unique capabilities in realizing lightweight designs with little2165penalty from geometrical complexity, and extensive demonstrations are available for various applications such as fashion, arts and biomedicine as shown in Fig.1 [1-5]. However, the design of these lightweight structures beyond aesthetic purposes is in general not well-understood by most designers. The lack of understanding on the relationship between various engineering performance requirements (e.g.

The lack of understanding on the relationship between various engineering performance requirements (e.g. mechanical properties, thermal properties, biological properties, etc.) and the geometrical design often prevents efficient design of lightweight structures for functional purposes. In addition, there also exist very little literature in the guidance of optimal process selections in the fabrication of these structures utilizing different AM technologies.a. Fashion shoes [6]b. Table [7]c. Skin shell [8]d. Artificial ear [9]Fig.1 Lightweight designs realized via AM in art, fashion and biomedicineWhile there exist various geometry design and optimization approaches and tools that generally allows for the creation of models with improved lightweight performance, currently none of them addresses the process designs adequately.

Most of the lightweight design tools treat material as an ideal isotropic material and focus on geometry optimization only, which often results in a significant design deviation from reality. One of the unique characteristics of AM is that the material properties are often process and geometry dependent. Such coupling effect has significant impact in the design practice of lightweight structures, since these structures often have geometrical features that have varying dimensions and therefore potentially varying material properties. In addition, due to the intrinsic quality variations with many AM processes with small-dimension geometries, the difficulty of achieving accurate design is further signified. In this paper some of the works in the attempt to tackle this design dilemma is presented.

In this paper some of the works in the attempt to tackle this design dilemma is presented. A design methodology that includes geometrical design and material property design was proposed, and some of the issues during the establishment of this methodology was discussed.Geometrical design of AM lightweight structures In the design of AM lightweight structures, two basic approaches are most commonly employed currently, which are topology optimization based design and cellular structure design. Each approach possess certain advantage and disadvantage compared to each other.Topology optimization is a mathematical optimization methodology that achieves an optimized design by redistributing material voxels within a given design space with the objectives of maximizing/minimizing certain criteria such as weight, stiffness, compliance and conductivity while being subjected to design constraints.

Unlike sizing optimization and shape optimization, topology optimization does not prescribe design topology, and is therefore capable of yielding at least locally optimized geometrical solutions [10]. During the optimization, individual voxels could be either removed, added or even partially filled2166when variable-densities are allowed [11-16]. The resulting structures often exhibit highly complex geometries that are challenging even for AM processes, and consequent re-design might be needed as shown in Fig.2. It was argued that despite the additional efforts needed, such design approach is still highly efficient in achieving optimized geometrical designs.CAD-baselineTopology optimizedRe-designedFabricatedFig.2 Topology optimization as design reference [17]Currently topology optimization design approach faces multiple challenges.

Among them are the lack of ability to handle certain optimization problems such as dynamic fatigue problems and multi-physics problems [18, 19], as well as the manufacturability issues. Besides the design limitations such as minimum feature sizes, orientation limits and support-free critical surfaces that could not be satisfactorily addressed due to the lack of definition with overall geometry, currently topology optimization also lacks the ability to handle anisotropic materials and materials with geometry-dependent material properties.Cellular structure designs, on the other hand, take a design approach with significant geometrical constraints in general. Traditional cellular structure design theories for stochastic cellular structures focus mostly on porosity designs.

Traditional cellular structure design theories for stochastic cellular structures focus mostly on porosity designs. Following Gibson-Ashby theory as shown in Eq.(1), various mechanical properties including elastic modulus (E), strength ((cid:305)) and shear modulus (G) can be determined via the relative density ((cid:545)r) of the cellular structures. In addition, various physical properties such as thermal transfer coefficient, coefficient of thermal expansion, electrical conductivity and mass transfer properties are also strongly dependent on (cid:545)r [20-24]. Following this approach, additional experimental verifications are usually needed in order to determine the scaling factors of individually manufactured cellular structures, which then allows for further structural designs using the characterized cellular structures as an equivalent “material”.

This design route is shown in Fig.3.E ρ∝2 r,σ ρ∝ 3/ 2 r,G ρ∝2 r(1)E E Sρ= 2 C 1 rG G Sρ= 2 C r 3σ σ S=Cρ 1.5 2 r2167Fig.3 Relative density based cellular structure designFor non-stochastic cellular designs that exhibit higher level of geometrical determinacy, although relative density based design approach is still applicable, addition design factors could be incorporated to enable more efficient multi-objective lightweight designs. When the cellular structures exhibit spatial geometrical periodicity, unit cell based design approach is often taken, which utilizes the geometrically representative unit cell to simplify the design of the entire structures. Compared to topology optimization, although limited by geometrical design flexibility, the unit cell design appears to show more promises in accounting for AM material properties during the design process due to the high predictability of geometries.

In this work, unit cell based cellular design was explored, and issues related to this design approach was discussed in details in the following section.Unit cell geometry design The design of periodic cellular structures with unit cell is analogous to the design of 3D tiling patterns. When only one type of unit cell geometry is used for the designs, regardless of the actual unit cell design, the geometrical bounding volume (GBV) of the unit cell must satisfy space-filling requirement. Fig.4 shows some typical space filling 3D elements that could potentially be used for cellular designs. Cubic GBV is currently most commonly utilized for unit cell designs. As shown in Fig.5, various unit cell geometries can be designed with cubic GBV. In general, these unit cell geometries possess good spatial symmetries and could therefore be further simplified during the design modeling.a. Cubeb. Hexagonal prismc. Triangular prismd. Octahedrone. Rhombic dodecahedronf.

Cubeb. Hexagonal prismc. Triangular prismd. Octahedrone. Rhombic dodecahedronf. Elongated dodecahedrong. gyrobifastigumFig.4 Typical space filling polyhedral [25]2168a. Octet-trussb. Octahedronc. Diamondd. RhombicFig.5 Several unit cell designs with cubic GBV One such example of modeling is shown below in details for the BCC lattice which is shown in Fig.6 [26]. The unit cell geometry for the cellular structure shown in Fig.6a could be represented by either Fig.6b or Fib.6c. However the unit cell layout 2 is easier to model.(a) lattice structure(b) Unit cell layout 1 Fig.6 BCC lattice cellular structure [26](c) Unit cell layout 2From the topology of the unit cell it is obvious that the uniaxial mechanical properties of the BCC lattice at all three principal directions could be modeled with identical formulations.

Considering a remote compressive stress (cid:305)z applied on the structure along the z direction, the loading of the unit cell is shown in Fig.7a. Since all struts are subjected to identical loading conditions and boundary conditions, an arbitrary strut is taken for modeling, whose loading condition is illustrated in Fig.7b. From force equilibrium the force components shown in Fig.7b can be determined as:F z1 4L Lσ=z1 2(2.1)M1=1 2F L z 3sin1 θ σ = z 8zL L L 1 2 3sinθ z(2.2)where L1, L2, are the dimensions of the unit cell in x and y directions, and L3 is the length of a half-strut, and (cid:537)z is the slope angle of the strut in relation to the z axis:θ z=sin− 12 L 1 2+L 32 L 2(2.3)2169(a) Loading of unit cell(b) Analysis of individual strutFig.7 Octahedral unit cell under compressionUnder the loading condition shown in Fig.7b, the strut undergoes a bending/shearing combined deformation.

Such conservative estimation could sometimes be justified by the fact that metal AM cellular structures often exhibit significant quality variability. From Fig.7b the normal stress and shear stress of the strut could be determined as:σ = 1M u F z+1 Icos Aθ z=σ zL L L 1 2 3 I 8sinθ zu+σ zL L 1 2 4cos Aθ z(8.1)τ = 1F sinz Ibθ zD(8.2)Where u is the distance from the geometrical center of the cross section to the location of interest, D is the moment of area of the cross section, and b is the width of the cross section at the location of interest.

Applying Von Mises Criterion, the maximum allowable stress level (cid:305)m that results in the onset of structure yield is:σ m=L L 1 22 L 32sin 2 I 2θ zu2+42 cos 2 A2θ z+θ z 2 2 I b26 sinD2σ Y(9)Following Eq.(7) and Eq.(9), through geometrical parameter design a range of mechanical properties could be achieved in particular direction of the BCC lattice structures. It is noted that although in Eq.(7) and Eq.(9) the material properties are constants, they could be readily substituted into material property functions, therefore potentially allows for multi- scale hierarchical designs with the cellular structures with material/process design factored incorporated.

Also, for cellular structures that could not be treated as beam networks, other existing solid structure modeling theories could be used, which further increases the flexibility of this design methodology.It must be noted that the underlying assumption for unit cell design approach is that the mechanical properties and responses of the unit cell are representative to the entire structures, which is only true when boundary conditions are ignored. However, for actual structures there always exist size effects. Size effects could be caused by both the lack of balancing forces at free boundaries and the boundary constraints as shown in Fig.8. Size effects vary from different unit cell designs and must be characterized either through experimentation or through modeling [28-30]. The size effects of several common AM cellular unit cell geometries, including the re-entrant auxetic structure, the octet-truss structure, the rhombic structure, and the BCC lattice structure, are shown in Fig.9.

Different unit cell designs exhibit different magnitudes of elastic modulus and maximum stresses when subjected to same amount of loading. In addition, the size effects along different directions of each unit cell geometries are also significantly different. Some unit cell designs exhibit highly predictable size effects that could be readily incorporated during the structural designs as correction factors, while the others require more studies in order to establish quantitative size effect design rules [28, 31].2171a. Free boundaryb. Boundary constraintsFig.8 Size effect mechanisms [31]a. Elastic modulusb. Maximum stress levels Fig.9 Size effects of several unit cell geometries2172Another assumption utilized in unit cell design that is often overlooked is the homogenization of cellular structures. It was often assumed that once the equivalent constitutive material properties of the cellular structures (i.e.

It was often assumed that once the equivalent constitutive material properties of the cellular structures (i.e. elastic modulus, shear modulus and Poisson’s ratios) are known, the cellular structures could then be effectively treated as a continuum for subsequent design. However, such homogenization treatment might need to be subjected to scrutiny. For example, for a regular hexagonal cellular structure shown in Fig.10, the two loading cases are equivalent for continuum according to principal stress rules when boundary conditions are ignored. If the cellular structures could be treated as continuum, then the two loading cases are approximately equivalent for the unit cell close to the center of the structure (highlighted in red box). However, it was observed that the differences of stress and deflection for a unit cell amount to about 25% and 45% respectively between the two cases.

However, it was observed that the differences of stress and deflection for a unit cell amount to about 25% and 45% respectively between the two cases. Although more studies are needed to further identify such effects, it was speculated that it might be largely caused by the semi- discontinuous and directional connectivity of the cellular structures. In addition, this also implies that the traditionally prevalent homogenization based methods might not always be as effective for cellular designs.a. Under normal + shear stress (cid:305)max=35.5MPa Unit cell distortion = 5.26x10-3mmb.

Under normal + shear stress (cid:305)max=35.5MPa Unit cell distortion = 5.26x10-3mmb. Under principal stress (cid:305)max=44.2MPa Unit cell distortion = 3.61x10-3mmFig.10 Comparison of cellular behavior under equivalent loading conditionsLightweight material design In general much less is understood with the manufacturability of the lightweight structures, although it has been identified that factors such as energy density, energy beam power, scanning speed, part location, part orientation and scanning strategies all have potentially significant effect on the mechanical properties of the cellular structures [32-36].Staircase effect tends to be more significant for cellular structures due to the small dimensions of the geometries. It has been suggested that at smaller orientation angles (i.e. more aligned to horizontal plane) the staircase effect of struts could become significant enough to affect the structural integrity as demonstrated in Fig.11a [37].

more aligned to horizontal plane) the staircase effect of struts could become significant enough to affect the structural integrity as demonstrated in Fig.11a [37]. As shown in Fig.11b-c, for thin struts fabricated by electron beam melting process, at 20° the cross section geometry of the strut exhibits more significant fluctuation compared to the 70° struts [38]. However, a recent work found contradictory trends for thin struts fabricated via laser melting process, in which lower-angle struts were found to be easier to fabricate due to their larger projected cross sectional areas [39].2173a. Staircase effect with different orientation angles [37]b. struts at 20° angle [38]c. struts at 70° angle [38]Fig.11 Staircase effect for cellular structuresAnother intrinsic effect that introduces geometrical error is the surface powder sintering for the powder bed fusion AM processes, which is caused by the heat dissipated away from the processed areas.

As shown in Fig.12, these surface defects causes dimensional variations on the cellular struts. For the calculation of mechanical properties, it was found that the minimum strut dimension (d in Fig.12b) should be used in the modeling [40]. On the other hand, in order to calculate pore size, the largest strut dimension (D in Fig.12b) should likely be adopted. For bulky structures such surface roughness could potentially reduce fatigue performance, and for cellular structures, this issue could be more pronounced due to the large specific surface areas of these structures. Literatures have shown that the fatigue strength of Ti6Al4V AM cellular structures are generally lower compared to that of the bulky Ti6Al4V fabricated via the same AM processes [32, 41-43]. Due to the complex geometries and the extensive existence of internal features, cellular structures are generally difficult to perform surface treatment with.

Due to the complex geometries and the extensive existence of internal features, cellular structures are generally difficult to perform surface treatment with. As a result, the fatigue performance of the metal AM cellular structures still poses a significant barrier for their structural applications.Dda. A Ti6Al4V strut from EBM processb. Dimensional variationFig.12 Surface quality issue with thin struts [40]Due to the geometrical errors, the actual dimensions of the cellular struts often deviate significantly from the designs, especially when the dimensions become smaller. Fig.13 shows some of the preliminary works with Ti6Al4V fabricated by EOS M270 laser melting system. In this experimental based works, thin struts with different orientations and dimensions were processed with the same beam energy (80W) and scanning speed (400mm/s) and different scanning strategies (contour + hatch and contour + edge) and beam offsets (0 and 40μm).

Significant dimensional errors occurred at all strut designs smaller than 0.5mm regardless of the orientation and other process conditions. It was also shown that when the beam offset was applied, the process was more capable of realizing2174lower-angle struts (15°) with higher accuracies. At dimensions larger than 0.5mm, the fabrication qualities of the struts become less sensitive to the selection of scanning strategies and beam offsets. Due to the intrinsic process variability, it was speculated that the random errors occur at smaller dimensions could not be effectively reduced. Therefore, such information should be used to provide lower threshold for the dimensional design of the struts in order to ensure the qualities.b. 45° struts Fig.13 Dimensional errors for Ti6Al4V thin struts fabricated by EOS M270a. 15° strutsc.

45° struts Fig.13 Dimensional errors for Ti6Al4V thin struts fabricated by EOS M270a. 15° strutsc. 75° strutsThe microstructure of the metal cellular structures fabricated via powder bed fusion AM often exhibits finer microstructrrual phases compared to the bulky structures. As shown in Fig.14, for Ti6Al4V cellular structures fabricated via electron beam and laser beam based powder bed fusion AM processes, the predominant microstructure is the fine-grained lath (cid:302)’ martensite, which is suggested to be contributed by the rapid cooling effect introduced by large surface area of the cellular geometries [44, 45]. In addition, it was observed that the size of the prior (cid:533) grains in the microstructure is also a function of both strut orientation and strut dimensions [39].

In addition, it was observed that the size of the prior (cid:533) grains in the microstructure is also a function of both strut orientation and strut dimensions [39]. As shown in Fig.15, for Ti6Al4V thin struts fabricated by laser melting process, with increasing feature dimensions, the microstructural grain size of the Ti6Al4V struts exhibit an increasing trend, while there also exist significant size differences between grains that are close to the exterior surface of the struts and those that are at the interior of the struts. On the other hand, smaller overhanging angle result in more consistent grain size distribution throughout the strut. Such microstructural variation likely corresponds to local mechanical property variations, which means that the mechanical properties of the cellular structures are likely coupled with their geometrical designs and process setup.a. Electron beam melting [44]b. Laser melting [45]Fig.14 Microstructure of AM Ti6Al4V cellular struts2175a.

Electron beam melting [44]b. Laser melting [45]Fig.14 Microstructure of AM Ti6Al4V cellular struts2175a. Interior grain sizeb. Exterior/boundary grain sizeFig.15 Grain size dependency on orientation and strut diameter for Ti6Al4V fabricated by laser melting [39]Very little works are currently available for the process optimization of cellular structures. In many metal powder bed fusion AM systems, the process parameters used for support structures are often more optimized towards cellular structures with very thin struts. However more systematic studies are needed. In addition, the scanning strategies could also become significant in determining the success of cellular structure fabrications. For example, it was observed that for the GP1 stainless steel fabricated with EOS M270, depending on the selection of scanning strategies, struts with certain dimensions might fail to be built as shown in Fig.16 [36].

This was contributed largely by the overlapping scanning paths of laser energy beams, which likely introduced certain types of defects into the structures and render the struts weak. The shearing effect of the powder spreader blade during powder recoating further signifies such defects and results in significant part distortion. For EOS M270 system, it was concluded that when contour + edge scanning were used, such critical dimensions exist and are closely associated with the overall beam offset [36].

0.2mm Fig.16 Thin strut fabrication using contour + edge scanning in EOS M270 [36]a. 0.1mmc. 0.3mm2176To summarize, the knowledge for AM cellular structure design is still rather fragmented now. A regularly employed shortcut is to perform experimental based structure characterization which, if planned properly, could result in efficient development of structural design guidelines for a particular cellular design. Various design factors such as process/material property relationships and size effects would be incorporated into the design process as compound effects through experimentation, and provide a short-term solution for companies that need design improvements.

On the other hand, a thorough understanding of different design aspects and the establishment of a systematic design methodology will benefit the design of cellular structures as a metamaterial for wider applications in the long term.Conclusions In this paper, the challenges of AM lightweight structure were briefly reviewed. Compared to other design methodology such as topology optimization, unit cell based cellular design functionality and to provide a good compromise between method appears manufacturability. However, in order to pursue this design approach, various additional factors must be considered, such as size effects and material property dependency on both geometrical designs of the struts and the process planning. It was found that the homogenization treatment could not be readily applied to cellular structures in general, which poses a rather challenging obstacle in adopting this design method in the design of actual structures.

Due to the complexity of cellular structures, it is currently inefficient to perform the designs using finite element simulation based methods. However, if the cellular structures could not be treated as continuous solid materials with equivalent properties, the limitation of analytical modeling must be overcome through over means.Acknowledgement This work was partially supported by Office of Naval Research (ONR) grant #N00014-16- 1-2394 and University of Louisville Intramural Research Initiation Grants. The authors would like to acknowledge the supports from Rapid Prototyping Center (RPC) at University of Louisville and from Center of Additive Manufacturing and Logistics (CAMAL) at North Carolina State University.Reference [1] I. Zein, D. W. Hutmacher, K. C. Tan, S. H. Teoh. Fused deposition modeling of novel scaffold architectures for tissue engineering applications. Biomaterials. 23(2002), 4: 1169-1185. [2] D. W. Rosen.

[5] B. Z. Wang, Y. Chen. The effect of 3D printing technology on the future fashion design and manufacturing. Applied Mechanics and Materials. 496-500(2014): 2687- 2691. Accessed July 2016. [6] http://www.3ders.org/articles/20120821-continuum-fashion-launches-custom-3d- printed-shoes.html. Accessed July 2016. [7] Francis Bitonti Studio. http://www.francisbitonti.com/fiber-tables/. Accessed July 2016. [8] Neri Oxman: Projects. http://www.materialecology.com/projects. Accessed July 2016. [9] Wake Forest Institute of Regenerative Medicine. http://www.wakehealth.edu/ WFIRM/. Accessed July 2016. [10] M. P. Bendsoe, O. Sigmund. Topology Optimizaiton: Theory, Methods and Applications. Springer-Verlag, Berlin, Germany, 2003. [11] Ing. T. Bechtold. Structural topology optimization for MEMS design. University of Freiburg, 2013. [12] G. Chahine, P. Smith, R. Kovacevic. Application of topology optimization in modern additive manufacturing.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDESIGN FOR INTERNAL LATTICE STRUCTURES WITH APPLICATION IN ADDITIVE MANUFACTURINGDonald Palomino1, Ryan McClelland2, Mike Grau3, Ryan Watkins4, Bingbing Li11Autonomy Research Center for STEAHM (ARCS), California State University Northridge, CA 91330 2Instrument Systems and Technology Division (ISTD), NASA Goddard Space Flight Center, Greenbelt, MD 20771 3Autodesk Inc, Novi, MI 94105 4Jet Propulsion Laboratory, California Institute of Technology, Pasadena, CA 91109AbstractInternal lattice structures have the potential to significantly reduce the mass of an existing metal component, which is a desirable characteristic in the aerospace and automobile industries. However, there are still uncertainties on whether or not internal lattice structures can outperform a solid version of the same mass.

However, there are still uncertainties on whether or not internal lattice structures can outperform a solid version of the same mass. Additionally, internal lattice structures can only be produced via additive manufacturing methods, bringing more challenges to resolve. To determine the viability of internal lattice structures, a study will be performed to compare its performance with solid, hollow, and mass penalty designs of equivalent masses using Autodesk Fusion 360. A performance baseline will be established by running multiple simulations on simple geometries to obtain the maximum displacement, first four modes, and first buckling mode.

A performance baseline will be established by running multiple simulations on simple geometries to obtain the maximum displacement, first four modes, and first buckling mode. A generative design part, better known within NASA Goddard Space Flight Center as A15, will undergo the same simulations and have its results analyzed to determine feasibility.Keywords: Metal Additive Manufacturing; Generative Design; Internal Lattice Structure1.IntroductionCAD modeling is a fundamental component in contemporary engineering and requiresboth skill and experience from the designer. Replacing the need for hand drafting on light tables, it has greatly increased the efficiency of the design process. However, this process itself has remained largely stagnant since. As the complexity of projects scale with time, so too does design work required. This is especially evident in the aerospace industry where the stiffness-to- mass ratio of parts is one of the most critical parameters.

This is especially evident in the aerospace industry where the stiffness-to- mass ratio of parts is one of the most critical parameters. One emerging technology to assist in this design phase is generative design.Generative design allows for an engineer to create multiple, valid solutions to a designproblem simultaneously and autonomously. This process can take what would take an experienced engineer weeks down to mere hours. To take advantage of this technology, a new methodology must be learned that contradicts the more traditional design process. However, the time invested into learning is more than made up in the increased productivity an engineer can achieve.1766Lattice structures, although a technology applied only recently, has actually been prevalent on Earth for millions of years. Everywhere from the hexagonal shape of honeycombs to the Voronoi structure that forms our bones, lattices are an essential building block within nature.

Everywhere from the hexagonal shape of honeycombs to the Voronoi structure that forms our bones, lattices are an essential building block within nature. Following this example, there exists different types of lattices, each optimized towards specific goals. Mimicking this design in modern design problems can provide multiple advantages, although manufacturing such complex designs remains a pressing issue.Internal lattice structures build upon this by adding a shell around the lattice, fully enclosing it. This has the potential to significantly increase the stiffness-to-mass ratio and buckling resistance for existing parts, yet is barely ever utilized in current industries because of a lack of research, education, and methods of validations for it. Additionally, they are fundamentally difficult to manufacture which further limits their use.In this paper, a method to effectively design for internal lattice structures for additivemanufacturing will be developed.

This will be accomplished by utilizing Fusion 360 by Autodesk and nTopology, both CAD softwares with internal lattice structure capabilities. A baseline will be established by running simulations on a simple geometry of three different designs: solid, hollow, and internal lattice. Each of these designs will be approximately the same mass to effectively compare them under strict mass budget scenarios, similar to that seen in the aerospace industry. This design process will also be well documented to provide a template on design for internal lattice structures to supplement future engineers and designers in potential applications. Lastly, a method to validate them once additively manufactured will be created. To be practical, it is important that these validation methods do not damage or destroy the part after manufacturing, also known as NDE (nondestructive evaluation).

To be practical, it is important that these validation methods do not damage or destroy the part after manufacturing, also known as NDE (nondestructive evaluation). Two methods will be looked into accomplishing this: using MRI (magnetic resonance imaging) and ultrasonic testing.1.1. Current Status of Metal AM, Generative Design, and Internal LatticeStructures1.1.1. Metal AMAdditive manufacturing (AM) grants incredible flexibility in manufacturing capabilitiescompared to traditional forms, like CNC machining. This process involves joining material, through various methods, together layer-by-layer to form one continuous part. Among its many advantages include reduction in material waste and logistics management1, mass optimization of parts, and complex geometry such as lattice and corrugation2. However, each layer’s stability is dependent on the prior layer.

However, each layer’s stability is dependent on the prior layer. If there is too great a difference in the current layer profile compared with the prior layer, then the overhanging material could sag and ultimately lead to a print failure3.To remedy this common problem, support structure can be added to an existing part toensure successful prints4. This is typically done autonomously by the 3D printing processing software, known as a slicer. Optimization of their ease of removal and reduction of material consumed is continuously being developed5. However, these support structures are only beneficial if they are able to be removed during post-processing. For internal lattice structures, this is not the case as the support material would be trapped inside without any method for1767removal. Therefore, a support-less AM method must be implemented to successfully print these complex structures6.

Therefore, a support-less AM method must be implemented to successfully print these complex structures6. Of the currently available options, two are the most applicable to this problem: powder bed fusion and direct energy deposition.Powder bed fusion (PBF) is a process where a laser beam points in a preprogrammedpath towards selected points of a flat bed of metal powder to fuse the material together. This is done either by full or partial melting, also known as selective laser melting (SLM) and selective laser sintering (SLS) respectively. To prevent oxidation of the metal and eliminate flammable gases like oxygen, the build chamber is vacuumed then filled with an inert gas7. This allows for a high XY resolution at the micron level, as well as having the capacity for very thin layer heights. Additionally, the powder bed acts as a support for the finished part, meaning no actual support structure is necessary. However, it does it come with its own challenges.

Additionally, the powder bed acts as a support for the finished part, meaning no actual support structure is necessary. However, it does it come with its own challenges. This includes various imperfections during the manufacturing process, such as surface roughness due to irregular powder size, porosity due to unmelted areas8, and perhaps most notoriously, trapped powder inside. This is perhaps the biggest limiting factor pertaining to internal lattice structures, as a majority of the complex geometry is enclosed in material. The conventional solution is adding a drain hole to the part to remove a majority of the powder, then heat treat is to sinter any residual powder left9.

The conventional solution is adding a drain hole to the part to remove a majority of the powder, then heat treat is to sinter any residual powder left9. This is not as simple in the case of internal lattice structures, as enclosed cells are likely to form and prohibit ease of powder removal.Direct energy deposition (DED) is a process that uses direct energy, typically in the formof a laser of electron beam, focused on a region to both melt the substrate and a deposited material onto the substrate, either a powder or wire. An important distinction between DED and PBF is the lack of a powder bed for DED, instead being more akin to a plasma welding machine. Similarly, it does not need support structures to produce complex shapes, but it has much poorer resolution and surface finish compared to PBF10.1.1.2.

Similarly, it does not need support structures to produce complex shapes, but it has much poorer resolution and surface finish compared to PBF10.1.1.2. Generative DesignGenerative Design (GD) is an iterative design process that uses artificial-intelligence based algorithms that involves a program that will generate a certain number of outputs that fulfill certain constraints. A designer is also involved in the process, selecting produced outputs or changing input values, ranges, and distribution. The designer does not necessarily have to be human; it can be an AI. The designer learns to refine the program with each iteration as their design goals become better defined over time11.This process can also be applied in great effect towards engineering application. Once aproject’s requirements have been properly identified, these can be input in a CAD software supporting GD, such as Fusion 360 by Autodesk.

Once aproject’s requirements have been properly identified, these can be input in a CAD software supporting GD, such as Fusion 360 by Autodesk. The outputs are designs that have been optimized for stiffness-to-mass ratio which is incredibly desirable in the aerospace industry. Typically, these results are organic in shape and require advanced manufacturing methods like 5- axis CNC machining and additive manufacturing, however it can be set to optimize its design for the relatively simpler 3-axis CNC machining12. This can significantly improve an engineer’s productivity as GD allows them to design parts with a higher stiffness-to-mass ratio in hours, whereas the same part may have taken weeks or months of time if done traditionally.

Additionally, adding manufacturing constraints increases the likelihood of a machine shop1768sending a quote, whereas a traditionally designed part may not receive any quotes at all due to difficulty of manufacturing13.1.1.3.Internal Lattice StructuresLattice structures, although a technology applied only recently, has actually been prevalent on Earth for millions of years. Everywhere from the hexagonal shape of honeycombs to the Voronoi structure that forms our bones, lattices are an essential building block within nature. Following this example, there exists different types of lattices, each specializing in a specific function such as mechanical, thermal, or even storage14. Mimicking this design in modern design problems can provide multiple advantages, although manufacturing such complex designs remains a pressing issue.Internal lattice structures build upon this by adding a shell around the lattice, fully enclosing it.

This has the potential to increase the stiffness-to-mass ratio, buckling resistance, and add vibration and noise dampening15. However, their complex geometry makes them difficult to manufacture, even with additive manufacturing. Multiple variables like printer resolution, material, and support structures can influence the success of these prints16. Additionally, internal lattice structures are difficult to model and simulate, severely limiting their methods for validation prior to manufacturing17.2. Methodology2.1. CAD Software & DesignFusion 360 by Autodesk is the CAD software used as it offers a suite of workspaces including modeling, generative design, and simulation. A beta feature called Alpha Centauri will be used to create internal lattices from inputted designs. Firstly, the Design Workspace is used to create two simple geometries: a straight cylinder and a curved cylinder. This is to set a baseline of reference for more advanced models to be compared to.

Firstly, the Design Workspace is used to create two simple geometries: a straight cylinder and a curved cylinder. This is to set a baseline of reference for more advanced models to be compared to. From this, four unique designs are applied: solid, hollow, internal lattice, and internal lattice with mass penalty. The first three designs are of the same mass for a fair comparison. This is done by first making the internal lattice design, then adjusting the diameter and shell thickness for the solid and hollow designs respectively until they reach the same mass as the former. The internal lattice with mass penalty features the same shell thickness as the hollow design, which is thicker than its equivalent mass internal lattice design.

The internal lattice with mass penalty features the same shell thickness as the hollow design, which is thicker than its equivalent mass internal lattice design. Once all four designs have been created for each simple geometry, it is run through a series of simulations to compare their performance against each other.2.2.Internal Lattice StructureTo design for internal lattice structures using Alpha Centauri, a solid body and internal volume body must first be created. The solid body is simply the model created from the Design Workspace. To produce the internal volume body, the solid body is copied and internally shelled. Then the shell body is subtracted from the copy to produce an internal volume representative of the defined shell thickness. These two bodies are inputted into Alpha Centauri, as well as the lattice type, cell size, and pipe diameter.

These two bodies are inputted into Alpha Centauri, as well as the lattice type, cell size, and pipe diameter. There are 15 unique lattice types available, however as1769of the time of this writing, it is not possible for a user to create their own unit cell. Cell size represents the size of the unit cell, and the pipe diameter is the thickness of the lattice struts. After processing these parameters, an internal lattice structure is produced and is ready to be simulated on.2.3. Generative DesignFigure 1: Setup for generative design with preserve geometries (green bodies) andobstacle geometries (red bodies)Once a baseline has been established with simple geometry, more advanced models will be developed using the Design and Generative Design Workspaces. The first step in doing so is setting up a design space consisting of bodies to later be defined as preserve and obstacle geometries in the Design Workspace.

The first step in doing so is setting up a design space consisting of bodies to later be defined as preserve and obstacle geometries in the Design Workspace. Unlike conventional part design, generative design only requires bodies that will be used as preserve or obstacle geometrics. Therefore, a starting design is not required. Once all bodies are created within the Design workspace, the rest of the setup process will continue in the Generative Design Workspace.Multiple inputs and parameters must be defined in the Generative Design workspacebefore a study can be successfully conducted. First is assigning the bodies made in the Design workspace as preserve or obstacle geometries. As their name implies, preserve geometries are assigned to bodies to incorporate them in the final shape of the design, while obstacle geometries are assigned to bodies to represent spaces for the design to avoid.

These geometries are assigned within the Generative Design workspace once the bodies in the Design workspace are finished. Figure 1 shows the preserve geometries assigned to the bodies representing the three bottom mounting pads and the large instrument interface at the top. Figure 1 also shows the obstacle geometry blocking material from generating in the instrument mounting location, the spacecraft face, and the bolt interfaces.1770Once all of the preserve and obstacle geometries have been assigned, design conditions including structural constraints and loads are defined. Structural constraints are to apply fixes, pins, or frictionless constraints to a geometry. In Figure 1 three pin icons can be seen on the bottom three mounting pads, indicating a pinned structural constraint. A pin is used instead of a fixed constraint as it is more accurate to how the part will be installed in the real-world using machine bolts.

A pin is used instead of a fixed constraint as it is more accurate to how the part will be installed in the real-world using machine bolts. Point masses are defined according to the center of gravity and mass of the part they represent, in this case being an optical instrument. These point masses must experience a G- load which is also applied according to the application. As these structures will undergo various G-loads during launch, this is also a more accurate way of modeling instead of subjecting faces to forces.Design criteria, such as design objectives, manufacturing method, and materials are alsodefined. Design objectives allow for the generative design study to prioritize minimizing mass or maximizing stiffness, as well as the factor of safety. Next is the manufacturing method, of which the following can be selected: unrestricted, additive, milling, 2-axis cutting, and die-casting. Finally, the material that the part will be manufactured from must be defined.

Finally, the material that the part will be manufactured from must be defined. While multiple materials can be selected to produce different results of their respective materials, Generative Design cannot apply these materials into a single result simultaneously.To give more volume for internal latticing to occur, an effort is made to alter theseparameters to force generative design to produce designs that are thicker and with no holes or struts. The most effective way to do so was found to be setting the manufacturing method to 3- axis milling, increasing the drill size diameter to 0.5 [in], and greatly increasing the target mass beyond the mission specification. Because of the larger drill size and the limitation of 3-axes, generative design produces parts that would normally be manufacturable given these constraints, resulting in non-complex bodies with generous filets. The high target mass allows the generative design to produce parts with ample thickness and volume.

The high target mass allows the generative design to produce parts with ample thickness and volume. Double, triple, and quadruple the initial target mass will be used to generate the thicker A15 brackets.Once all of the results have been generated, filters are applied to select the stiffest parts,and to select the parts with the highest mass for hollow and latticed parts. These are then exported as part files where their organic body can be modified to make thin areas thicker and avoid self-intersecting geometry, a major obstacle in the workflow to producing internal lattice structures.2.4. Avoiding Self-Intersecting GeometryPerhaps the most time-consuming challenge in producing internal lattice structures are the self-intersecting geometries that are produced in Fusion 360. This is due to the simulation’s inability to properly mesh the design if these are present.

This is due to the simulation’s inability to properly mesh the design if these are present. The software does not currently have a tool or feature that automatically resolves them, instead any corrections necessary must be done manually. It has been found that almost all self-intersecting geometry occurs where material generated closest to the bolt interface preserve geometries, as the material surrounding the obstacle geometry is thin and overlaps with each other when thickening the organic body.1771The above steps in Section 3.4. to producing parts through generative design helps toreduce self-intersecting geometries, however they will likely still appear even with preventative steps. To remedy these occurrences, the following, additional steps should be taken. Firstly is to grab the organic body near the bolt interfaces and pull the faces apart to prevent overlaps from developing once it is thickened.

Firstly is to grab the organic body near the bolt interfaces and pull the faces apart to prevent overlaps from developing once it is thickened. This has to be done while making sure that the organic still overlaps the preserve geometry, otherwise a combine failure will occur instead. Even after this step, it is still likely that self-intersecting geometries will develop, so a last step to solve this is to use the section analysis tool and comb through the design to find them. Once found, an arbitrary shape should be made that covers up the region of self-intersection. Using a series of combine tools, this shape will fill in the self-intersecting geometry similar to filling in a cavity in a tooth.

Using a series of combine tools, this shape will fill in the self-intersecting geometry similar to filling in a cavity in a tooth. Once all of the self-intersecting geometries have been resolved, the part will properly mesh, and simulations can occur as normal.17722.5.SimulationFigure 2: Static Stress Simulation of Lattice Straight Cylinder with Load Applied Relative to Perpendicular of Top Surface: a) 0 [deg], b) 45 [deg], c) 90 [deg]After the parts produced using generative design have been modified appropriately andall self-intersecting geometries have been resolved, they can be brought to the Simulation Workspace. From here, three simulations are conducted to acquire the necessary data of the part to assess their performance: a Modal Frequency simulation for the first four modes, a Static Stress simulation for the maximum displacement and stress, and a Buckling simulation for the first buckling mode.

An example of the static stress simulation for the lattice straight cylinder is shown in Figure 2 where blue regions represent areas of low stress and red regions represent areas of high stress. Each simulation was set up by fixing the bottom face and applying a 100 [N] load to the top face. This load starts perpendicular to the applied face at 0 [deg], then is incremented by 15 [deg] until reaching 90 [deg], parallel to the applied face. The resultant deformation in each subfigure is adjusted to better show the effect of the applied force.For A15 brackets, Fusion 360 can automatically take the setup parameters used forgenerative design to populate the necessary inputs for conducting simulations. The same three aforementioned simulations are run. The only major difference is the inclusion of steel bolts in the top instrument interface to best represent realistic conditions.

The same three aforementioned simulations are run. The only major difference is the inclusion of steel bolts in the top instrument interface to best represent realistic conditions. Otherwise, these holes may collapse on themselves, which is incorrect as the steel bolts would prevent this.17733.Results3.1.Simple Geometry ResultsFigure 5: a) Solid Straight Cylinder, b) Hollow Straight Cylinder, c) Lattice Straight Cylinder, d) Lattice Straight Cylinder with Mass Penalty, e) Solid Curved Cylinder, f) Hollow Curved Cylinder, g) Lattice Curved Cylinder, h) Lattice Curved Cylinder with Mass PenaltyIn total, eight simple geometries were designed using Fusion 360 as shown in Figure 5.These represent three different designs for testing: solid, hollow, and internal lattice. Each design for their respective cylinder was made to be approximately the same mass to ensure a fair comparison.

Each design for their respective cylinder was made to be approximately the same mass to ensure a fair comparison. Therefore, while the length for each cylinder remained as 100 [mm], their diameters and shell thicknesses vary where applicable. A fourth model called internal lattice with mass penalty was also made to provide results for an internal lattice cylinder with the same shell thickness as the hollow cylinder. This is because the shell thickness of the internal lattice cylinder is slightly less than that of the hollow one to reach similar masses. The unit cell of the lattice was 25 [mm], the beam diameter was 5 [mm], and the beam thickness was 0.5 [mm]. Due to current limitations with Alpha Centauri, these were the parameters that produced the most consistent results.

Due to current limitations with Alpha Centauri, these were the parameters that produced the most consistent results. Smaller parameters would result in errors.1774Figure 6: Performance Results of Straight CylinderAfter finishing the simulations, the results are graphed against each other to compare thesolid, hollow, internal lattice, and internal lattice with mass penalty variations. From this, four graphs are produced as shown in Figure 6. The maximum displacement at 90 [deg] of the solid variation compared to the internal lattice with mass penalty variation is 86.5% more. All four variations are considerably resistant to buckling based on the applied loads. The stiffness-to- mass ratio at 90 [deg] of the solid variation compared to the internal lattice with mass penalty variation is 836.04% more.

The stiffness-to- mass ratio at 90 [deg] of the solid variation compared to the internal lattice with mass penalty variation is 836.04% more. For the first modal frequency of the solid variation compared to the hollow variation is 140.3% more.1775Figure 7: Performance Results of Curved CylinderSimilar results and trends as the straight cylinder are shown for the curved cylinder inFigure 7. The maximum displacement at 90 [deg] of the solid variation compared to the internal lattice with mass penalty variation is 86.8% more. All four variations are considerably resistant to buckling based on the applied loads. The stiffness-to-mass ratio at 90 [deg] of the solid variation compared to the internal lattice with mass penalty variation is 856.1% more. For the first modal frequency of the solid variation compared to the internal lattice with mass penalty variation is 139.6% more.17763.2.

For the first modal frequency of the solid variation compared to the internal lattice with mass penalty variation is 139.6% more.17763.2. Unit Cell ResultsFigure 8: Unit Cell Lattices a) Pyritohedron, b) Dualtx, c) Vin Tiles, d) SoftboxAs one of the purposes of this paper is to determine the most optimal design for internallattice structures, another series of simulations were run on curved cylinders with different lattice structures as shown in Figure 8. Different unit cells may afford better performance despite having the same shell thickness and body. Also, some may be easier to produce via additive manufacturing than others.

Different unit cells may afford better performance despite having the same shell thickness and body. Also, some may be easier to produce via additive manufacturing than others. For example larger overhang angles, measured from the horizon, are beneficial towards successful, un-supported prints.1777Figure 9: Performance Results of Curved Cylinders with Different Lattice StructuresUsing the four, unit cells in Figure 8 with the internal lattice with mass penalty curvedcylinder design, the resulting graphs in Figure 9 are produced. Since the shell thickness and body is the same with the four cases, the results are very similar to one another. For maximum displacement, the Softbox lattice experiences 4.93% more displacement than the Vin Tiles at 90 [deg]. Similarly, Vin Tiles perform marginally better than the other lattice types for buckling mode and stiffness-to-mass ratio.

Similarly, Vin Tiles perform marginally better than the other lattice types for buckling mode and stiffness-to-mass ratio. Vin Tiles also performed best for Modes 2 and 4, but overall the differences in modes were slim with only a 1.81% difference between the Pyritohedron and DualTX designs for Mode 1.3.3. A15 ResultsFigure 10: A15 Variations a) Solid, b) 400g Target Mass, c) 600g Target Mass, d) 800g Target MassOnce the baseline has been established with the results obtained from the simple geometry designs, the same series of versions and simulations are done to the A15 bracket, as shown in Figure 10. To ensure sufficient internal volume for hollowing and latticing to occur,1778the A15 is set up to be thick and without thin beams, unlike Figure 10a). Both hollow and internal lattice structures are formed from Figure 10b), c). and d).Figure 11: Internal Lattice vs.

Both hollow and internal lattice structures are formed from Figure 10b), c). and d).Figure 11: Internal Lattice vs. Hollow: a) Internal Lattice from 400g Target Mass, b) Internal Lattice from 600g Target Mass, c) Internal Lattice from 800g Target Mass, d) Hollow from 400g Target Mass, e) Hollow from 600g Target Mass, f) Hollow from 800g Target MassA section view of the thicker A15 versions for both hollow and internal lattice structuresare shown in Figure 11. To ensure all variations are approximately the same mass, the hollow variations have a shell thickness of 1.5 [mm], 1.3 [mm], and 1.1 [mm] respectively to Figure 11a), b) and c).

To ensure all variations are approximately the same mass, the hollow variations have a shell thickness of 1.5 [mm], 1.3 [mm], and 1.1 [mm] respectively to Figure 11a), b) and c). For the internal lattice variations, all three’s shell thickness is 1 [mm] with an appropriate lattice density to match the mass of the solid variation.1779Figure 12: Performance Results of Solid, Hollow, and Internal Lattice A15 VariantsUsing the same three simulations in Fusion 360 used for the simple geometries (staticstress, modal frequency, and structural buckling), the performance of the solid, hollow, and internal lattice variants of A15 were obtained, as shown in Figure 12. The average maximum displacement of the internal lattice and hollow A15 variants is 0.0254 [mm] and 0.0207 [mm] respectively, an 18.5% decrease.

The average maximum displacement of the internal lattice and hollow A15 variants is 0.0254 [mm] and 0.0207 [mm] respectively, an 18.5% decrease. The average first modal frequency of the internal lattice and hollow A15 variants is 147.5 [Hz] and 149.2 [Hz] respectively, a 1.2% increase.4.Discussion4.1.Simple GeometryThe simple geometry forms a baseline to observe resultant patterns from the data. Asseen in Figure 6 for the results of a straight cylinder, the maximum displacement, stiffness-to- mass ratio, and first modal frequency of the hollow and internal lattice variants are overall considerably better than the solid variant, with the hollow and lattice variants showing closely related trends. The hollow variant produced 30.0% less displacement, 19.4% greater stiffness-to- mass ratio, and 9.4% greater first modal frequency compared to the lattice variant.

The hollow variant produced 30.0% less displacement, 19.4% greater stiffness-to- mass ratio, and 9.4% greater first modal frequency compared to the lattice variant. When compared to the lattice with mass penalty variant, these results change with the hollow variant having 5.7% more displacement, 34.3% lesser stiffness-to-mass ratio, and 7.1% greater first modal frequency. This shows that the hollow variant is overall the better performing design for the simple hollow cylinder.Figure 7 shows the results for a simple curved cylinder. Similar to the results obtained from the straight cylinder, the maximum displacement and first modal frequency of the hollow and internal lattice variants are overall considerably better than the solid variant, with the hollow1780and lattice variants showing closely related trends. However, the stiffness-to-mass ratio of the lattice with mass penalty variant is noticeably better compared to the hollow and lattice variants.

However, the stiffness-to-mass ratio of the lattice with mass penalty variant is noticeably better compared to the hollow and lattice variants. The hollow variant has an 81.5% greater stiffness-to-mass ratio compared to the solid variant, but the lattice with mass penalty variant has a further 77.3% increase in stiffness-to-mass ratio. However, the hollow variant has only a 3.7% advantage over the lattice variant. As the lattice with mass penalty variant is the same as the lattice variant but with a greater shell thickness, this increase in stiffness-to-mass ratio may be attributed more due to the shell thickness rather than the internal lattice structure. Putting it aside, the hollow and lattice variant behave very similarly across the board.4.2. Unit CellOne of the biggest advantages in additive manufacturing is free complexity, which iswhere the manufacturing cost per piece does not increase relative to increase in geometric complexity18.

This gives merit to determining the best performing unit cell for the internal lattice structure. Based on the results shown in Figure 9, this was found to be the Vin Tiles unit cell as it consistently had the lowest maximum displacement and a high first modal frequency. Although it only yielded less than 5% better properties compared to the worst performing unit cell, the free complexity afforded by additive manufacturing makes this small improvement worth the effort.4.3. A15As one of the major issues with producing internal lattice structures and hollow designsin Fusion 360 is the generation of self-intersecting geometries, three different A15 brackets were made based on increasing target masses inputted during generative design. Additionally, the three hollow variants based on these generative design results were made with slightly increased shell thicknesses to approximate the mass of the solid and lattice variants.

Additionally, the three hollow variants based on these generative design results were made with slightly increased shell thicknesses to approximate the mass of the solid and lattice variants. Based on the results shown in Figure 12, there does not seem to be a clear trend forming between the hollow or lattice triplets. The 400 [g] and 600 [g] target mass lattice variants perform worse than the solid A15, but the 800 [g] target mass variant performs 25.9% better than the solid variant. This sudden increase in performance may be due to the wider diameter of the A15 due to the high initial target mass. An example of this can be found in nature through the baobab tree in Africa which features an incredibly wide trunk diameter with a very low inner density. It was found that increasing trunk diameter directly increases its strength against buckling19. Similarly, the 600 [g] and 800 [g] target mass hollow variants perform best in their group.

It was found that increasing trunk diameter directly increases its strength against buckling19. Similarly, the 600 [g] and 800 [g] target mass hollow variants perform best in their group. This may be explained by the wider diameter compared to the solid and 400 [g] target mass hollow variant. However, when looking at the first modal frequency of all seven A15 variants, there is only a 14.3% difference between the worst performing variant, solid, and the best performing one, 800 [g] target mass lattice. While this is indeed an improvement, it may not be considerable enough to justify the additional time necessary towards producing such a design, at least considering the current methodology. Lastly is the stiffness-to-mass ratio, from which the 800 [g] target mass lattice variant and hollow group has the highest ratio. However, as seen in Figure 11, there is very little infill at all in the 800 [g] target mass lattice variant.

However, as seen in Figure 11, there is very little infill at all in the 800 [g] target mass lattice variant. This suggests that a high stiffness-to-mass ratio is more dependent on the diameter width, similar to a baobab tree. Therefore, internal lattice structures do not offer a significant enough advantage over solid designs to offset the associated risks and difficulties of additive manufacturing. However, hollow designs may have the potential1781to outperform solid designs if the diameter width and shell thickness are optimized. To this point, it can be seen with the solid A15 variant that it closely resembles a hollow structure. The center of the part is completely void of material, with the outer geometries seemingly wrapping around it. These outer geometries likely show the maximum load paths that the part experiences. All of the empty spaces in between them can be thought of as zero thickness walls.

These outer geometries likely show the maximum load paths that the part experiences. All of the empty spaces in between them can be thought of as zero thickness walls. As these spaces in the center of the part and between the outer geometries do not experience high loads or stresses, no material is required to support the structure. This shows that the generative design process converges on hollow designs as the most optimal stiffness-to-mass parts.4.3.1. Wider Diameter A15 VariantsFigure 13: Wider Diameter A15 Variants: a) Hollow (2kg TM), b) Section View, c) Hollow ("Brick"), d) Section ViewBased on the new hypothesis that hollow structures with a wide diameter will result in stiffer parts, two new designs were produced using generative design. As seen in Figure 13a) and b), these parts were produced by inputting a target mass of 2 [kg] and taking an early iteration from generative design respectively.

As seen in Figure 13a) and b), these parts were produced by inputting a target mass of 2 [kg] and taking an early iteration from generative design respectively. By adjusting the shell thickness, both parts are1782approximately 140 [g] like the other A15 variants. The results of these two variants are listed below in Table 1 to contrast with the already established variants. It should be noted that due to the high volume that these parts have, proper care should be taken to ensure that there is sufficient space in their installation location.4.4.

This supports the hypothesis that wider diameters in hollow structures result in higher stiffness. The best of these results comes from the “brick” hollow variant which was obtained by rolling back the iterations of a generative design result to reveal a design that was as thick as possible. Conversely, parts with a lower target mass during generative design resulted in higher maximum displacements, lower modal frequencies, and lower stiffness-to-mass ratios. Therefore, it can be stated that wider diameters directly correlate to higher stiffness for hollow structures. Additionally, internal lattice structures do not offer greater stiffness compared to hollow designs.5.ConclusionDespite recent advancements in software and additive manufacturing technologies allowing for the modeling, simulation, and fabrication of internal lattice structures, the A15 case study does not yet offer a clear advantage compared to contemporary designs.

While this may not be universally indicative of all applications, it does point to this conclusion for similarly behaving parts used in the aerospace industry. Parts produced from generative design have1783flexibility in being optimized for 3-axis & 5-axis CNC machining and additive manufacturing, however, internal lattice structures can only be produced via the latter. With this comes additive manufacturing’s restrictions in build volume, non-homogeneous material, and difficulty in validation. Similar to internal lattice structures, hollow structures can only be produced by additive manufacturing, however they are much easier to model, simulate, and post-process. Both hollow and internal lattice structures require drain holes to remove excess powder after printing, however, it is more difficult to remove said powder from the latter than the former.

Both hollow and internal lattice structures require drain holes to remove excess powder after printing, however, it is more difficult to remove said powder from the latter than the former. Fusion 360 is not able to automate internal lattice structures, requiring more time to be expended by the designer or engineer to develop. Lastly, based on the simulation results of the multiple A15 variants, internal lattice structures do not offer significantly better performance compared to the much more feasible hollow structure. Therefore, if additive manufacturing is an option, efforts should be directed toward producing hollow structure to achieve the highest stiffness-to- mass parts with minimal work needed.For future work, efforts should be diverted toward shell thickness optimization and increasing diameter width of the part.

The results of this paper have shown that diameter width is one of the driving variables in decreasing maximum displacement, which is one of the most important performance metrics considered for aerospace applications. Additionally, by optimizing shell thickness, more material can be dedicated to high stress areas while conserving material in areas of low stress. This could both save powder usage and further minimize mass. Additionally, another path of research worth considering is variable density hollow structures. This takes inspiration from naturally occurring hollow structures like bamboo and baobab trees, both of which increase fiber density towards their outer edge where stresses are the highest19 20. Optimizing the density of sintered metal powder in the structure could further increase stiffness in conjunction with shell thickness optimization.1784References[1]Khajavi, Siavash H., et al.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesign for protection: Systematic approach to prevent product piracy during product development using AMU. Jahnke*, R. Koch*, A. T. Oppermann*Direct Manufacturing Research Center (DMRC) and Chair of Computer Application and Integration in Design and Planning (C.I.K.), Paderborn University, Mersinweg 3, 33098 Paderborn, GermanyAbstractAlthough infringements of intellectual properties in terms of product piracy are growing for years and threaten investments in research and development most companies still rely on legal measures like property rights. A more preventive effect to protect against counterfeits can be achieved using technical measures complicating reverse engineering, improving traceability and assuring data protection.

A more preventive effect to protect against counterfeits can be achieved using technical measures complicating reverse engineering, improving traceability and assuring data protection. Additive Manufacturing can contribute a lot to the effectivity and efficiency of those technical measures but presently they are often unconsidered during product development. To support decision makers and designers through all the steps of a product development process an integrated systematic approach has been developed. Protective measures using AM are allocated to specific process steps and responsible persons in charge so that the result is a guideline for “design for protection”.

Protective measures using AM are allocated to specific process steps and responsible persons in charge so that the result is a guideline for “design for protection”. The main idea is to help developing piracy-robust products for that the return of investment is not threatened by counterfeits and its economical impacts.Introduction and motivationBranches dealing with mechanical and electrical engineering and manufacturing of systems like the automotive or electronics industry currently benefit a lot from globalization, decreasing costs for logistics and advanced infrastructures of information and communication technologies (ICT). [AKL11] Based on innovation and unprecedented productivity due to automation of manufacturing systems the German machine and plant engineering industry came up with up to now highest turnover volume of 218.1 billion € in 2015.

[VDMA16] Between 2012 and 2015 more than every second company in mechanical engineering developed new products and/or processes and gained more than 20% or their turnover with novelties on the market. To continue on this success companies spend billions of € in their capabilities for innovations. [ZEW16] Exactly those investments and generated competitive advantages are highly threatened by imitations. PORTER stated already in 1985 out that the sustainability of competitive advantages can only be kept by implementation of inherent barriers hampering rivals to imitate innovation independent from the legal point of view. [Por85] As globalization, decreasing costs in logistics and the proceeding ICT infrastructure pushes innovations on the one side it brings up new risks on the other side. [WAB+07] [BMF16] In relation to simple consumer goods the number of imitations of highly complex products up to whole machines is increasing further.

[WAB+07] [BMF16] In relation to simple consumer goods the number of imitations of highly complex products up to whole machines is increasing further. [VDM16]2481It seems like the words of Confucius are still very newsworthy:"By three methods we may learn wisdom: First, by reflection, which is noblest; Second, by imitation, which is easiest; and third by experience, which is the bitterest.“Unfortunately it is the second methods that is used more and more. Regarding the initiative “business action to stop counterfeiting and piracy (BASCAP)” the easiest method of learning wisdom ends up in economical damages of roughly 1.770 trillion US $ only in 2015. [BAS11] More detailed information are shown in biennial studies by the German Engineering Association (VDMA). According to their study “Produktpiraterie” 70% of the associated companies are affected by imitations with a total damage of 7.3 billion € in 2016.

According to their study “Produktpiraterie” 70% of the associated companies are affected by imitations with a total damage of 7.3 billion € in 2016. Additionally and very hard to quantify are damages that will occur in the future in terms of loss of images caused by imitation not satisfying or even endangering customers. Particularly the risks for customers’ health, especially for children’s health, are focused in the European H2020 project “iBUS – integrated business models for customer driven custom supply chains”. Innovative solutions to customize toys and therefore come up with added values and increased satisfaction are in development in this project. To meet European regulations regarding toys’ safety is one the most important topic when thinking about customer driven designs. Imitations, not only poor and low quality copies, are a considerable threat violating European regulations in many cases.

Imitations, not only poor and low quality copies, are a considerable threat violating European regulations in many cases. Derived from the findings in the iBUS project and the alerting figures shown in the studies a preventive protection against product piracy and counterfeits should be integrated in innovation and development processes of companies across different branches. At least the awareness has to be raised to make a conscious decision. [VDM16]Today most companies try to protect their innovations and knowledge with legal measures that are mostly reactive except of a deterrent effect. But as the study made by the VDMA shows the most important source for product pirates is the reverse engineering re- generating product data from the physical product. Legal measure will not be able to prevent reverse engineering as the illegal activity is very obvious in most cases.

Legal measure will not be able to prevent reverse engineering as the illegal activity is very obvious in most cases. Furthermore law is different according to the territorial principle and it is very hard to consider all relevant countries. [Lor12] In contract to that technical measure have an effect that is much more protecting against reverse engineering and not restricted regionally but only 29% of questioned companies implement technical measure in their products. Most of them because of their felt assessment that there are no technical measures available and effective.Additive Manufacturing with its freedom in design, flexibility in production and those often mentioned benefits is able to contribute with new technical measures preventing product piracy. The main limitation using additive manufacturing and its protective technical measures is that a product’s design has to be manufacturable additively.

The main limitation using additive manufacturing and its protective technical measures is that a product’s design has to be manufacturable additively. The fewest products designed for conventional manufacturing like injection molding or milling can be produced economically and/or technically in additive manufacturing processes without changing the design. Therefore the main intention of this paper is the integration of additive manufacturing and related technical possibilities to prevent imitations in standard development processes to support companies and designers to design for protection. According to KOKOSCHKA a lack of knowledge and awareness in the companies is responsible for the now enormous need of action. [Kok13] Thus the systematic approach will take all departments into account so that the whole topic will be considered company-widely in the future.

[Kok13] Thus the systematic approach will take all departments into account so that the whole topic will be considered company-widely in the future. A holistic, sustainable and integrated knowledge management will be a solution to spread the potentials of additive manufacturing, its applications to protect products against product piracy and recommendations for actions. In a first step the systematic will be embedded in a paper catalogue to be used by all players involved in the development of innovations.2482Application centered requirements analysisDivided in two subsections, formal and application based requirements, this section deals with the analysis and definition of requirements to be fulfilled by the systematic approach.

The main aim is to consider all demands and issues so that the methodology will be integrable or applicable in the companies’ process map.Formal requirementsThe objective of this subsection is to ensure a systematic and structured development and definition of the systematic approach respectively methodology. Thus the main requirements here are based on PATZAK [Pat82] who stated out that methodological models should be meet on the following five requirements to be effective: Empirical correctness: Model and empirical study will arise similar results,  Formal correctness in terms of reproducibility and traceability,  Case related to achieve a solution or benefit for the focused problem,  Usability for easy use of model as well as for interpretation of results and  Low-effort-applicability so that effort and profit have a good relationFurthermore rules of systems engineering are considerable as the systematic approach will technical systems.

support protective development of complex products up HABERFELLNER summarizes those rules in four main thoughts: [HDWFV12]to Development from rough to detailed  Thinking in and about variations  Breakdown in phases  Enabling of cyclic problemsolvingApplication based requirementsRequirements shown in table 1 are resulting out a of more deep analysis of background in terms of where does the problem come from and how are companies acting in the development of new and innovative products. The table is divided in three sections of requirements on the systematic approach: Regrading product protection, regarding knowledge management and regarding additive manufacturing:Product protection requirements are shown from R1 to R7.

Knowledge management requirements to be fulfilled by the systematic approach are specified in R8 and R9 and finally as the whole approach aims at enabling companies to use additive manufacturing for a protection against counterfeits R10 brings that into account.2483Table 1: Requirements for systematic approach in three sections: Regrading product protection, regarding knowledge management and regarding additive manufacturing2484Analysis of systematic approaches for product developmentThere are a couple of approaches to foster the protection of product and therefore to increase the resilience against product piracy. Without going too much into detail all of those approaches have been assessed regarding their degree of fulfillment the previously defined requirements. Detailed analysis of each approach has built the basis for the rating shown in table 2.

Detailed analysis of each approach has built the basis for the rating shown in table 2. All necessary information needed for transparency and reproducibility of the assessment are available in the primary sources given in the following section. The existing approaches have been clustered in three sections regarding their main aim: Development of product-protection strategiesPreviously to the research offensive driven by the Federal Ministry for Education and Research (BMBF) in Germany in 2010 there have been three approaches existing that are wide spread in the literature made byJACOBS et al., FUCHS and VON WELSER und GONZALES. [JSJ01][FKMR06][WG07] In addition to those there came up a new one driven by the offensive published by GASSMANN und BECKENBAUER in 2010. Imitation-protected design of products [GB10]Four further approaches focusing a more application based procedure have also been developed during projects funded by the BMBF since 2010.

Imitation-protected design of products [GB10]Four further approaches focusing a more application based procedure have also been developed during projects funded by the BMBF since 2010. The main objective here is to directly design protected product so that the strategic point of view is mostly missing. In the following table these approaches are named MEIMANN and the mentioned projects with their acronyms KoPiKomp, ProAuthent and KoPira. [Mei10][BJSR10][GDS10][AAG10] Development of imitation-protected product and processesLast but not least the considered approaches in this category combined already the good parts of both categories before so that they became more holistic. In comparison to the approaches mentioned before these ones are considering the areas procurement, manufacturing, distribution and information management as well. The considered approaches are described in detail by NEEMANN, SCHNAPPAUF, ABELE, MEIWALD, KLEINE und KOKOSCHKA.

The considered approaches are described in detail by NEEMANN, SCHNAPPAUF, ABELE, MEIWALD, KLEINE und KOKOSCHKA. [Nee07][Sch10][AKL11] [Mei11][Kle10][Kok13]Finally table 2 and the assessed approached respecting the defined requirements are showing a very white spot in consideration of AM potentials and in applicability as a holistic approach during the whole development process. None of the analyzed methodologies supports from the very first idea on a strategic layer over product development up to the development of the production system. Exactly this consistency through the whole process chain is missing resulting in coordination effort afterwards. Considering requirements and potentials of the production system directly at the beginning will help to avoid expensive changes prior to the start of production.

Considering requirements and potentials of the production system directly at the beginning will help to avoid expensive changes prior to the start of production. Therefore table 2 shows the need for action having the main objective in mind: Bringing additive manufacturing technologies more into application increasing the image regarding product piracy as it is often seen as a technology pushing imitations and counterfeits.2485Table 2: Assessment of currently existing approaches for development of protection strategies, protected design of products and imitation-protected products and processes2486Modelling of design for protection systematic approachAs a first step the main idea to model a systematic for a design for protection approach is to go along with standard design methodologies that are wide spread and really used in the industry.

Therefore from a German perspective the VDI2221 is used as a basis and will be modified in some phases to fulfill the requirements regarding the integration and consideration of additive manufacturing as well as knowledge management.

Phases, task and methods to be applied as well as the phases’ results are described in figure 1 that is divided in two sections again: The first phase in light grey has only to be carried out once for a company in a longer time period while the other phases in dark grey have to be carried out for each product to be protected.Figure 1: Assessment of currently existing approaches for development of protection strategies, protected design of products and imitation-protected products and processes2487Integration of systematic in standard product development processThe integration the design systematic shown before in standard product development processes will enforce designers as well as product managers from a more strategic point of view to think about measures to prevent product piracy endangering for example in the toy industry children’s safety and health.

Also in other industries and branches product piracy might always come along with less safety due to inferior materials or manufacturing processes as the main aim here is to generate profit.A model divided in three cycles focusing on the strategic product planning, the product development itself as well as the productionsystem development is one of the most holistic methodologies used in industrial and scientific areas. [GPW09] The main aim is to give an overview about the steps and responsible persons that are affected by the systematic approach. This will support an easy integration in companies’ everyday life. One advantages of the three-cyclic-model acc. to GAUSEMEIER et al. is the interdisciplinary applicability so that it can be used for each product independent from the branch or industry up to smart products with a high degree of automation and intelligence.

is the interdisciplinary applicability so that it can be used for each product independent from the branch or industry up to smart products with a high degree of automation and intelligence. [GEW09] Figure 2 shows the adapted model and assigned activities.Figure 2: Integration of systematic approach for a design for protection using AM in the three-cyclic- model from business planning to start of production acc. to GAUSEMEIER [GEW09]2488Validation of systematic approach protecting an industrial applicationTo validate the designed systematic approach to protect products against product piracy and counterfeits and further misuse an industrial application and its transformation from a very first idea to a protected additively manufactures product is presented in table 3.

All the step defined in figure 1 have been carried out with resulting in a very resilient product.Table 3: Assessment of currently existing approaches for development of protection strategies, protected design of products and imitation-protected products and processesApplication example: Sample collector system for bioreactorsFigure 3a: First idea for convential manufacturing Brief descriptionFigure 3b: Part „designed for protection“ with AMThe shown sample holder carousel is designed for the use in a sample collecting system to be used for automatic sampling of nutrient solution of bioreactors. The first idea (on the left side) shows a design for conventional manufacturing fulfilling most of the intended functions but without any protection against counterfeiting. Proofing potential risks in the 5 step approach the resilience against reverse engineering has been assessed to be very poor. Following to that this part has been selected during the identification phase.

Following to that this part has been selected during the identification phase. It is very important for the intended functionality of the whole system as well as manufacturable additively. Thinking about potential protective measures already published at the SFF a couple of measures has been selected with very good synergies. [JBR15] The measures and resulting benefits are described in the following cell. Finally the last phase aimed at realization of the analyzed functional structure considering the selected measures during the design.

Finally the last phase aimed at realization of the analyzed functional structure considering the selected measures during the design. In terms of knowledge management this table can act as a kind of fact sheet to be used by designers to benefit from already achieved success stories.Manufacturing process:Selectives Lasermelting (SLM)Material:AlSi10MgBatch size:depending on demand - envisaged: 100+/yearApplied AM potentials and resulting benefitsTool-less manufacturing: Due to the manufacturing using SLM in this application example conventional manufacturing has only been used for removing of support materials.Flexible Production: During development there have been six iterations for draft designs until the final one has been proofed to be robust manufacturable and protected against reverse engineering. As there are no tools needed the effort was very low to go through these2489itereations. Only changes in 3D CAD has been necessary.

As there are no tools needed the effort was very low to go through these2489itereations. Only changes in 3D CAD has been necessary. After development phase this potential is very useful for the economical production driven by demand.Freedom of design: Using the possibilities in design that are offered by AM a functional integration has been realized resulting in a design for protection.Individualization: Each part is marked with an individual code for traceability so that counterfeits can be distinguished from the original one as there are visible and invisible markings as well.

Therefore an identification becomes possible on the one hand and authentication on the other.Safety against product piracy: All the mentioned potentials applied in this design for protection end up in safety against product piracy as for example reverse engineering is complicating by far and traceability is pushed.Conclusion and outlookSummarizing the systematic approach and its validation safety against product piracy can be achieved using this structured methodology. Due to the fact that the VDI2221 design methodology is very wide spread and in industrial application the defined approach is very usable as there are a just modifications made to integrate missing topics like AM conformity and knowledge management. In this paper knowledge management is coming off badly as the whole approach need to be used over a longer term so that a kind of design catalogue can arise showing more example like the sample collector.

In that way designer will be supported learning from already made success stories. This is exactly what needs to happen in the future. The process steps and phases are spread in the industry and there will come back feedback and applications to be collected in supporting guidelines to foster more and more designs for protection.AcknowledgementsThe research leading to these results has received funding from the Federal Ministry for Education and Research (BMBF) in Germany in the project “it’s OWL 3P – Prevention of Product Piracy“. The text reflects the authors' views. BMBF is not liable for any use that may be made of the information contained therein. For further information see http:// http://www.itsowl-3p.de/The research leading to these results is co-funded by the Horizon 2020 Framework Programme of the European Union under Contract No. 646167 (Project iBUS). The text reflects the authors' views.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDesign for the Additive Manufacturing Process ChainDavid W. Rosen The G. W. Woodruff School of Mechanical Engineering Georgia Institute of Technology Atlanta, GA 30332 e-mail: david.rosen@me.gatech.eduDigital Manufacturing and Design Centre Singapore University of Technology and Design Singapore, 487372AbstractPost-processing operations are required for most additive manufacturing (AM) processes. For production parts, consideration of these post-processing operations during design is critical to achieve design requirements. For both metal and polymer parts, the sequence of steps in the process chain can be extensive. A design framework called the Process Chain Map (PCM) is introduced in this paper that explicitly relates design requirements for the part to each step in the AM process chain.

A design framework called the Process Chain Map (PCM) is introduced in this paper that explicitly relates design requirements for the part to each step in the AM process chain. This PCM visually shows the role of each step in the process chain and facilitates communication among design and manufacturing personnel. Software implementation of the PCM enables generation of system-level problem formulations of multidisciplinary design optimization problems. An example of a metal AM part demonstrates the PCM and the formulation of such a design problem.Keywords: design for additive manufacturing, manufacturing process chain, process chain mapIntroductionPost-processing operations are required for most additive manufacturing (AM) processes. Consideration of the entire process chain, from the AM process to all the post-processing operations, during design is critical to achieve design requirements.

Consideration of the entire process chain, from the AM process to all the post-processing operations, during design is critical to achieve design requirements. For metal parts, heat treatment, support structure removal, and finish machining are frequently required to convert the AM-fabricated part to its final state. Also, polymer parts usually require some post-processing operations, including support removal and finishing [1]. A technique called the Process Chain Map (PCM) is introduced in this paper that explicitly relates design requirements for the part to each step in the AM process chain. This PCM visually shows the role of each step in the process chain and facilitates communication among design and manufacturing personnel.A framework for part design across the process chain is presented based on the PCM. The objective of the framework is to ensure that part designs are optimized for performance and for the capabilities and limitations of the manufacturing processes.

The objective of the framework is to ensure that part designs are optimized for performance and for the capabilities and limitations of the manufacturing processes. Accordingly, each step in the process chain should be optimized for the system-level objectives, in addition to part performance optimization. Leveraging multidisciplinary design optimization methodology, the framework includes design optimization problem formulation where the system level design problem is decomposed into a series of subsystem problems that correspond to the steps in the process chain. Requirements embodied in the PCM are incorporated into the design problem53formulation. Influences of manufacturing processes on each other throughout the process chain are modeled using coupling variables among the corresponding subsystems.An example of a metal AM part demonstrates the PCM and the formulation of a design for the process chain problem.

The specific part is a bike stem to be fabricated in metal powder bed fusion (PBF). Results show that the PCM successfully models relationships between design requirements and the steps in the process chain and guides the designer in developing part designs for the process chain.Process Chain MapThe process chain map (PCM) is intended to be a visual representation of the relationships between steps in the process chain and how those steps contribute to satisfying design requirements. An example PCM for metal PBF is shown in Figure 1. Across the top is the sequence of manufacturing processes in the process chain. The first column contains the design requirements for the part in terms of shape, tolerance and surface finish requirements, and mechanical properties. The user of the PCM needs to decide the level of detail of the requirements. For a specific part design, it may be important to include different tolerances if they are achieved by different manufacturing processes.

For a specific part design, it may be important to include different tolerances if they are achieved by different manufacturing processes. In this example, different mechanical properties are included for bulk regions of the part, as well as surface properties, but this may not be relevant to all parts. In the body of the PCM, entries (green checks) indicate which processes contribute to achieving the requirement.In this illustration, the process-requirements relationships are shown qualitatively. An alternative usage of the PCM is to include quantitative values in the table to indicate the extent to which the requirement is achieved or the specific values of, for example, mechanical properties achieved by the process. This is similar to the development of Quality Function Deployment charts [2] that capture qualitative relationships between customer requirements and engineering characteristics.

This is similar to the development of Quality Function Deployment charts [2] that capture qualitative relationships between customer requirements and engineering characteristics. More complete implementations of QFD added quantitative entries so that the importances of engineering characteristics could be rank-ordered.Figure 1: Process chain map for metal powder bed fusion.The PCM can serve multiple purposes. First, it documents the process chain for the part so that all participants in product development can understand the sequence of processes. The PCM aids communication among the product development team; if someone does not understand the need for a process, they can discuss with the relevant manufacturing personnel.

The PCM aids communication among the product development team; if someone does not understand the need for a process, they can discuss with the relevant manufacturing personnel. Also, the PCM clarifies the role of each process in the process chain, indicating to which requirement(s) it contributes.54Further, software implementations of the PCM can be used for various design, manufacturing, and management purposes.• Relate design requirements to specific processes in the process chain • Visual depiction aids understanding of each process’s role • Communication aidLay out the entire process chainAdditive Manufacturing Process ChainsAs illustrated in Fig. 1, AM process chains can be extensive.

1, AM process chains can be extensive. Final part characteristics can be rather different than those produced by the AM process and it is imperative that designers design with final part properties and characteristics.In metal PBF, parts are fabricated with support structures, also called anchors, that fix the part to the build platform and prevent the part from warping due to residual stresses. Support structures also conduct heat from the laser to the build platform to reduce the likelihood of hot spots. These supports need to be removed by a machining operation: cutting, EDM, machining, etc. However, if supports are removed while the part has significant residual stresses, it can deform considerably. To prevent this, a stress relief heat treatment is typically performed while the part is still attached to the build platform.

To prevent this, a stress relief heat treatment is typically performed while the part is still attached to the build platform. The stress relief not only relaxes residual stresses, but can also modify and homogenize metal microstructures and significantly reduce mechanical property anisotropies. Additionally, annealing, solution annealing, and aging heat treatments may also be performed to modify microstructures and properties, in addition to (or instead of) stress relief.After PBF, parts are buried in the powder bed. It is necessary to remove the loose powder. Sometimes, people use brushes to remove powder particles from part surfaces, crevices, or holes. Automated depowdering stations are also available to automatically rotate build platform with parts to facilitate powder removal from different directions and from internal channels.

Automated depowdering stations are also available to automatically rotate build platform with parts to facilitate powder removal from different directions and from internal channels. Part surfaces tend to be fairly rough after metal PBF so various surface finishing operations may be performed on parts before or after removal from the build platform. Blasting with beads, sand, and ice is common. Finish machining is also common for part surfaces that have high accuracy, shape, or surface finish requirements.Most of this discussion applies to metal directed energy deposition (DED) processes as well, except that parts are not buried in a powder bed. Parts are typically built on a build platform; parts either need to be removed from the platform, or the platform is incorporated into the part design, but often needs to be cut to shape. Heat treatments and finish machining are common.For polymer processes, the process chain is also important.

Heat treatments and finish machining are common.For polymer processes, the process chain is also important. Support structures are used in most polymer processes, with the exception of PBF. Surface treatments (bead blasting, sanding) are commonly used to provide desired surface finish. Some machining or hand work may be performed to ensure correct shapes and accuracy requirements are met.In summary, parts are rarely ready for use after only the AM process. One or more post-processing steps is used frequently to ensure parts satisfy requirements. Some of this post-processing can significantly impact part characteristics and properties so it is important to capture their effects.Design for the Process Chain MethodA framework has been developed for the design of parts for the process chain, based on methods from Multidisciplinary Design Optimization (MDO) [3,4].

The objective is to design part shapes and find dimension values that satisfy as well as possible design requirements and constraints. Many times these55requirements and constraints will not be satisfied by the AM process, which is why subsequent post- processing steps are needed.The proposed approach starts with the formulation of the overall design problem, that is referred to as the part-process chain co-design problem. An example problem formulation is shown in Fig. 2 (based on [5]). At the top, the overall problem formulation is given as finding part shapes and dimensions, part characteristics, and process settings for all of the manufacturing processes in the process chain. This MDO problem seeks to minimize a measure of deviation from goals on, for example, performance requirements (e.g., maximum stress, maximum deflection), economic objectives (build time, cost), and weight.

Design variables, constraints, and objectives are often related to the requirements listed in the process chain map.The overall problem is decomposed into various subsystems that represent each step of the process chain from the process chain map, as well as part performance design. The problem formulation is illustrated with subsystem design optimization problems, but the overall design method need not distribute optimization in this manner, which will be discussed later. An advantage of this approach is that typically different analyses are performed for the different subsystems and these analyses can be performed independently from one another. The disadvantage is that the various subsystem design problems cannot be solved independently. This is easy to see if we consider mechanical properties. Elastic modulus and yield strength, among other properties, are needed for structural performance analysis of the part being designed.

This is easy to see if we consider mechanical properties. Elastic modulus and yield strength, among other properties, are needed for structural performance analysis of the part being designed. If fatigue is important, then surface roughness is important to consider. Since these properties and characteristics are the results of the entire process chain, the subsystem design problems are coupled, which has important implications. First, the problem coupling implies that some quantities are coupled, which may include design variables or other intermediate variables. Mechanical properties, such as elastic modulus, are good examples of coupled quantities. Second, the problem structures usually determine the sequence that subsystem problems should be solved.

Mechanical properties, such as elastic modulus, are good examples of coupled quantities. Second, the problem structures usually determine the sequence that subsystem problems should be solved. With few exceptions, the subsystem solution sequence should be the same as the process chain, with part design being the last subsystem to be solved.Figure 2: Product-process chain co-design problem formulation.To deal with subsystem coupling and identify suitable work flows, we use an extended design structure matrix diagram [6], as shown in Fig. 3. In the middle part of the figure, the four boxes denote the four subsystems from Fig. 2, but note that part design corresponds to the box labelled ‘D4 – Design.’ D1-D3 are56the three manufacturing processes in the process chain. Information flows are indicated by the green solid arrows that connect the boxes. Output variables from D1 flow to D2, D3, and D4. Similarly, output variables from D2 flow to D3 and D4, etc.

Information flows are indicated by the green solid arrows that connect the boxes. Output variables from D1 flow to D2, D3, and D4. Similarly, output variables from D2 flow to D3 and D4, etc. Additionally, outputs from D2-D4 can flow back up to D1 and potentially the other boxes (the diagram is shown fully connected, but need not be so). Solid black arrows pointing downwards from the D1-D4 boxes indicate optimization objectives, J1 to J4 respectively, that are computed by the subsystems. These objective quantities are inputs to the system-level performance evaluation where they are used to compute overall optimization objective function J*. Execution proceeds sequentially from D1 to D4 then to the system level performance evaluation.

Execution proceeds sequentially from D1 to D4 then to the system level performance evaluation. After evaluation, the system level optimization tests for convergence; if not converged, iterations continue starting again at D1.With this presentation, it should be clear that the process chain map can be used to directly develop part design optimization problems for the process chain that can be solved using MDO methods.Figure 3: Information and execution flows in the MDO problem expressed using an extended design structure matrix diagram.ExampleA bike stem part will be used to illustrate the application of the process chain map and concepts around design for the process chain. A bike stem attaches the handlebars to the steering fork as illustrated in Fig. 4. We assume that bike stems are loaded in bending, torsion, and in tension.

A bike stem attaches the handlebars to the steering fork as illustrated in Fig. 4. We assume that bike stems are loaded in bending, torsion, and in tension. The titanium alloy Ti-6Al-4V is selected for the material and we will plan for production (high volume) manufacturing using PBF.Requirements for the bike stem include supporting those expected loads, light weight, cylindricity tolerances on the cylindrical bores (large holes that mate with the steering shaft and the handlebars), a perpendicularity tolerance on the two bores, and surface finish requirements overall (< 40m Ra), in the bores (<m Ra), and on the sides of the bores (< 1m Ra).57Figure 4: Bike stem part.Topology optimization will be used to reduce part weight. The AM process chain consists of the PBF process followed by a stress relief heat treatment and an aging heat treatment to ensure maximum mechanical properties are achieved. Support structures will be removed using EDM.

Support structures will be removed using EDM. Parts will be bead blasted to ensure that surface finishes overall are below 40m Ra. Finally, finish machining will be performed on the bores and side surfaces to ensure that tolerances and surface finish requirements are met.Figure 5: Process chain map for the bike stem part.As can be seen in the right-most column of Fig. 5, all requirements can be achieved using this process chain map. Since the engineering work has not yet been completed, however, this conclusion should be considered as preliminary and an indication of what should be possible. The role of each step in the process chain can be understood by identifying requirements to which each contributes. For example, the interpretation of the stress relieve process is that it relieves residual stresses which enables the overall stem to maintain its desired shape and ensures the bosses (for the bolts and nuts used for attachment) have correct shapes and positions.

Aging contributes to mechanical properties but does not have much impact on part or feature shape. Bead blasting affects overall surface finish and contributes to surface mechanical properties58(surface treatments typically induce compressive stresses). Machining fine-tunes feature shapes and enables achievement of tolerances and surface finish requirements.Results to date are as follows. Topology optimization results are shown in Figure 6a. After further design, the final part is shown in Fig. 6b. To maximize the number of parts that fit on a build platform, the stems were oriented vertically; lattice structure struts were adjusted so that they could be fabricated without requiring support structures. Bosses were added for the bolts and nuts required to fasten the stem. After finalizing the part design, the PBF build layout and support structures were designed. A view of the stem oriented vertically is included in Fig. 7a.

After finalizing the part design, the PBF build layout and support structures were designed. A view of the stem oriented vertically is included in Fig. 7a. Some support structures are needed in the cylindrical bores. It was decided to use a teardrop shape for these bores, effectively adding gusset supports in the bores, instead of conventional vertical support structures since gussets should lead to faster machining times. Orthogonal views of the stem are shown in Fig. 7b. The bosses on the sides of the top bore require support structures, as can be seen in Fig. 7c, which shows an array of 20 bike stems on a single EOS M290 build platform. Solid supports were used at the bottom of the stem to facilitate heat conduction into the build plate, as can be seen in Fig. 7d.topology optimized designb) final designFigure 6: Bike stem designs.Process planning of machining operations was performed to remove support structures, machine the bores, and slit the bores to permit assembly.

Simulations of these operations were also conducted, the results of which are shown in Figure 8.The design for the process chain framework has not been fully implemented yet, so results from the application to this example are not yet available.ConclusionPost-processing operations are required for most additive manufacturing (AM) processes. As a consequence, consideration of these post-processing operations during design is critical to achieve design requirements, hence the need for design for the AM process chain methods. The Process Chain Map (PCM) was introduced in this paper as a technique to explicitly relate design requirements for the part to each step in the AM process chain. This PCM visually shows the role of each step in the process chain and facilitates communication among design and manufacturing personnel. A framework for design for the process chain was introduced based on methods from MDO.

A framework for design for the process chain was introduced based on methods from MDO. An example of a metal bike stem part demonstrates the PCM and the formulation of such a design problem. In conclusion, the PCM appears to be a useful technique for documenting and communicating manufacturing process chains. Further, the PCM can assist in formulating design for the process chain multiobjective optimization problems.59Figure 7: Bike stem PBF manufacturing: a) single stem positioned on build plate, b) two views of bike stem showing teardrop hole shape in bores, c) array of bike stems on build platform, d) support structures at base of part.Figure 8: Machining process planning with estimates of machining times.60AcknowledgementsPartial support from the Advanced Manufacturing and Engineering, Industry Alignment Fund program in the Singapore National Research Foundation, grant A19E10097, is gratefully acknowledged.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDESIGN FOR(E!) ADDITIVE MANUFACTURING: IN SEARCH OF A COMPREHENSIVE DESIGN CHALLENGE SUITABLE ACROSS AM EDUCATIONNicholas Meisel School of Engineering Design and Innovation The Pennsylvania State University, State College, PA 16802AbstractModern engineering design education relies heavily on the concept of problem-based learning (PBL). Driven by the constructivist theory of education, PBL enables students to build knowledge organically, rather than through rote memorization. As such, design for additive manufacturing (DfAM) education also tends to emphasize the use of PBL to encourage student learning. Unfortunately, dedicated DfAM education is still nascent. The result is a wide range of educators leveraging an equally wide, and often unproven, range of design challenges to support DfAM PBL.

Unfortunately, dedicated DfAM education is still nascent. The result is a wide range of educators leveraging an equally wide, and often unproven, range of design challenges to support DfAM PBL. Because of this, there is the possibility that a chosen design challenge will not represent AM as a true end-use manufacturing process nor promote a design space that can benefit from the full consideration of all opportunistic and restrictive DfAM concepts. In this paper, the author discusses the creation and implementation of a comprehensive design challenge that is suitable across the range of AM education. Specifically, the author proposes the use of a golf putter DfAM design challenge. This concept draws from lessons learned over years of DfAM instruction at undergraduate and graduate levels and is based in the need for three key aspects for a successful DfAM challenge in education: (1) clarity, (2) applicability, and (3) demonstrability.1.

IntroductionModern engineering design education relies heavily on the concept of problem-based learning (PBL) [1]. Based in the educational framework of constructivism, PBL leverages an organic approach to learning, where students naturally construct understanding through contextualized, hands-on learning [2]. This framework aligns naturally with Design for Additive Manufacturing (DfAM), which requires engineers to reconsider existing rulesets surrounding what makes a traditional design “manufacturable.” Through PBL, engineering students can engage with DfAM in a way that allows them to tangibly challenge their existing design intuition. Especially with the rapid advancement of desktop material extrusion in the classroom, it is becoming easier and easier to engage students with AM and DfAM in the classroom.However, even though the use of AM in the classroom continues to grow quickly, DfAM education is still fractured.

Most commonly, DfAM forms one portion or module of a more general AM course [3]. To support this module, many courses will leverage a design challenge, where students are tasked with designing or redesigning an object with the specific intent of using AM to manufacture the solution. This mirrors the growing popularity of internet-based DfAM challenges, such as the well-known GE bracket challenge [4]. Unfortunately, chosen DfAM PBL challenges are widely varied, which can introduce a range of inconsistencies in AM education across institutions and educational levels. The result is that students, upon entering2137industry, may not all demonstrate the same command of DfAM knowledge, since they have constructed their DfAM knowledge through different contexts.

The result is that students, upon entering2137industry, may not all demonstrate the same command of DfAM knowledge, since they have constructed their DfAM knowledge through different contexts. This commands the need for the DfAM education community to coalesce around a common design challenge, capable of meeting the best-practices associated with design PBL, while also demanding students consider the entirety of AM’s design opportunities and restrictions. This paper takes the first steps toward proposing such a unifying design challenge driven by the key characteristics of (1) clarity, (2) applicability, and (3) demonstrability. The result is an argument for the golf putter as a potential candidate for DfAM educational challenges.2. Background2.1. The Challenge of Addressing Dual DfAM in the Classroom Before creating a consistent approach to teaching DfAM in higher education, it is first necessary to establish a shared mental model about what defines DfAM.

One approach is through a dichotomous view of DfAM separated into ‘opportunistic’ and ‘restrictive.’ In this model, the DfAM principles that leverage AM capabilities can be collectively thought of as ‘opportunistic’ DfAM, which consist of (1) mass customization [5,6], (2) part consolidation [7] and printed assemblies [8], (3) free shape complexity [9–11], (4) embedding external components [12], and (5) printing with multiple materials [13]. Meanwhile, design guidelines meant to account for AM limitations and minimize failure [14] can be considered as ‘restrictive’ DfAM. This encompasses (1) the need for support structures [15], (2) warping of parts due to thermal stresses [16], (3) anisotropy and weakness in build direction [17,18], (4) surface roughness due to stair-stepping [19,20], and (5) limited feature size and accuracy [21].

When considered simultaneously, restrictive and opportunistic DfAM come together to form a paradigm of ‘dual DfAM,’ which can help support the integration of all DfAM concepts into the engineering design process [22– 29]. Table 1 summarizes some of these key DfAM considerations, divided into restrictive and opportunistic categories.Table 1. Key Restrictive and Opportunistic DfAM ConsiderationsRestrictive DfAM Support Structure Accommodation Warping Due to Thermal Stresses Delamination and Material Anisotropy Stair-Stepping and Surface Roughness Minimum Feature SizeOpportunistic DfAM Geometric and Hierarchical Complexity Multi-Material Printing Part Consolidation and Printed Assemblies Mass Customization Functional Complexity and EmbeddingEven after identifying the concept of dual DfAM as framework for AM education, there is still a need to identify how best to appropriately integrate it within engineering education.

It is not enough for a DfAM educational intervention to introduce students to different AM processes and DfAM concepts; it must also naturally encourage students to integrate both opportunistic and restrictive DfAM into their design process [30]. Fortunately, a range of DfAM initiatives have been introduced into the engineering curriculum, as summarized in [31]. However, research shows that if students are not motivated to fully explore the solution space enabled by AM [32], they may still generate simple, easy to manufacture geometries despite being trained in opportunistic DfAM [33]. Students may also focus only on avoiding build failure, which could lead them to apply only restrictive DfAM and not fully leverage opportunistic DfAM. This is problematic though, since sufficient emphasis on both opportunistic and restrictive DfAM is2138necessary for students to generate the most high-quality design concepts for AM [34].

This is problematic though, since sufficient emphasis on both opportunistic and restrictive DfAM is2138necessary for students to generate the most high-quality design concepts for AM [34]. Because of this, it is essential to understand how different design challenges can encourage students to engage with the design process while motivating them to fully leverage dual-DfAM [35].2.2. Existing Efforts in Problem-Based Learning for DfAM Education As discussed, research has shown that task-based learning helps students develop their problem solving capabilities [36], which have been identified as crucial for the continued growth of successful AM engineers [34]. Because of this, a key characteristic of DfAM education is that it encourages students to actively use AM and DfAM concepts to solve problems [37,38]. Task- based learning techniques can help achieve this.

Task- based learning techniques can help achieve this. One of the first instances of DfAM task-based education focused on technical and economic viability of introducing AM as a prototyping tool in engineering education [39]. More extensive use of PBL to solve a DfAM task was demonstrated in [3], where students were tasked with designing an AM solution for a problem of their own choosing. A similar, though more extensive use of PBL was also demonstrated through the implementation of a university-wide vehicle design competition, where students were tasked to create printable ground and aerial vehicles capable of maneuvering across obstacles while carrying a defined payload [35].Even beyond these initial examples, the use of PBL across AM and DfAM education is extensive.

This includes researchers combining short design activities focused on restrictive AM (geometric accuracy, strength, resolution, and process parameters) as well as longer design activities focused on opportunistic DfAM (cellular structures and lightweighting) [40]. In this particular case, students were encouraged to leverage DfAM concepts to design a car bumper capable of minimizing damage in the event of a collision with a wall. PBL for DfAM education is also demonstrated in [41], where students were asked to identify a problem of interest and generate solutions to solve it; information about AM and DfAM was presented to students as their worked through the design process to help students adapt their thinking to DfAM.

Longer- form DfAM workshops have also leveraged PBL methods, such as the redesign of a block manifold across four days in [42] and the redesign of a vehicular air breather in [43].Though this review is non-exhaustive, it provides a representative cross-section of the range of design challenges present in PBL across DfAM education. Often, design challenges are (1) student-identified and inconsistently scoped based on personal interests, (2) instructor identified ‘toy’ problems with clear bounds but limited design opportunities, or (3) industry-driven redesign tasks with vague bounds, but significant design opportunities. However, the author argues that successful PBL in DfAM requires a design challenge at the intersection of these three options. A task is needed with opportunities for personal design creativity, but with clear design bounds for an end-use industry product.

A task is needed with opportunities for personal design creativity, but with clear design bounds for an end-use industry product. Even further, the task must encourage students to consider all aspects of both restrictive and opportunistic DfAM, rather than just a limited number of concepts (such as a lightweighting task which might only require consideration of lattice structures or shape complexity).3. The AM Golf Putter Design ChallengeConsidering the current state of the art for PBL in DfAM, there is a need for a unifying design challenge that can engage a range of students from across educational levels with the totality of2139dual DfAM concepts toward a true, end-use product.

To that end, this section details an argument for the use of a golf putter design challenge to support DfAM based on the fundamental principles of clarity (whereby the task is grounded in a familiar problem with digestible design requirements), applicability (whereby the task benefits from full consideration of all aspects of restrictive and opportunistic DfAM), and demonstrability (whereby the task results in an end-use product that can be created and tested in a classroom environment).3.1. Clarity: Familiar Problem with Digestible Design Requirements The first characteristic of a successful DfAM challenge is the notion of clarity. Clarity is based primarily on the idea that the chosen task is grounded in concrete, real-world product design, with established objectives and constraints. Further, the task should be easily digestible to a range of students with a variety of backgrounds.Common familiarity/access to real-world experience.

Further, the task should be easily digestible to a range of students with a variety of backgrounds.Common familiarity/access to real-world experience. Since the onset of the COVID-19 pandemic, participation in golf has grown at a rate not seen in 17 years [44]. This is especially relevant for universities, as many have affiliated golf courses for student use. Despite this, it is still possible, and likely, that students have not participated in a full golf course. However, the ubiquity of mini golf as a recreational activity drastically increases the likelihood that students have hands on experience with golf putters; annual participation in mini golf involves 18 million players [45]. This familiarity with the challenge’s context naturally primes students for participation.Clear, concrete design constraints. The governing committee behind golf in the United States, the USGA, has a freely available document detailing all design constraints associated with golf putters [46].

The governing committee behind golf in the United States, the USGA, has a freely available document detailing all design constraints associated with golf putters [46]. Example constraints are shown in Figure 1. This gives students easy access to a series of concrete rules driving the size, weight, and form of compliant golf putters. This includes specifics regarding how AM’s “free complexity” may ultimately support or conflict with established rules, such as the requirements that clubs be “plain in shape.” Crucially, though there are a series of guidelines, they are not so limiting as to make the solution space trivial.Figure 1. Example Guidelines as Provided by the USGA [46]Performance driven by singular engineering objective. In putter design, modern products are typically driven by strict consideration of the putter’s moment of inertia (MOI). By increasing the MOI, the putter is generally considered to be more stable.

In putter design, modern products are typically driven by strict consideration of the putter’s moment of inertia (MOI). By increasing the MOI, the putter is generally considered to be more stable. The result is that the club face is less likely to be unintentionally opened or closed when swinging to impact the ball. The goal of increasing moment of inertia is understandable for engineering students early in their education and is typically intuitively addressed by the notion of moving weight to the perimeter of the putter.2140Abundance of existing, varied design inspiration. Due to the constant advancement of golf technology and design, there is an abundance of design inspiration available to students to help encourage their creativity and understanding of the task. Putters are typically categorized as blade, half-mallet, and mallet designs, all of which approach the concept of moment of inertia and usability differently.

Putters are typically categorized as blade, half-mallet, and mallet designs, all of which approach the concept of moment of inertia and usability differently. Industry is also starting to explore the use of AM in putter design, though efforts are still nascent. Currently, Cobra Golf is the only company with a commercially available putter with AM components, as seen in Figure 2. The prevalence of such numerous putter designs enables students to quickly understand the potential breadth of the solution space.Figure 2. Commercially Available Putter with AM Components (Adapted from [47])3.2. Applicability: Benefits from All Aspects of Opportunistic and Restrictive DfAM The second characteristic of a successful DfAM challenge is the notion of applicability. As discussed in Section 2, many existing design challenges tailored to DfAM benefit from only a select number of DfAM considerations, most frequently geometric complexity and support material elimination.

By contrast, the design of a golf putter has the potential to beneficially leverage the entire spectrum of restrictive and opportunistic DfAM. A non-exhaustive list of such examples follows in the rest of this sub-section, considering opportunistic DfAM first, followed by restrictive DfAM.O-DfAM 1: In-situ embedding. Though typically not as common as other aspects of opportunistic DfAM in product design, the putter challenge naturally encourages students to consider the use of in-situ embedding as a DfAM opportunity. This is especially true if their printer access is limited to common desktop-scale material extrusion systems. Such systems are unable to generate products with sufficient mass capable of matching traditional putter designs (approximately 350g).

Such systems are unable to generate products with sufficient mass capable of matching traditional putter designs (approximately 350g). As such, students tend to quickly recognize the need to embed additional weights into the putter.O-DfAM 2: Shape complexity and topology optimization: As discussed in Section 3.1, increasing the MOI of the putter head can improve stability during the swing and reduce undesired twisting. By leveraging DfAM’s shape complexity, when combined with topology optimization or generative design, it is possible to distribute material toward the perimeter of the club head to achieve a lightweight, yet stable putter in a way that is not traditionally possible with conventional manufacturing.2141O-DfAM 3: Hierarchical complexity and lattices. As with shape complexity, hierarchical complexity and lattice structures can be used to redistribute material to increase the MOI of the putter.

As with shape complexity, hierarchical complexity and lattice structures can be used to redistribute material to increase the MOI of the putter. However, they can also be used to provide texture to the face of the putter to change the feel on impact, as well as change the way in which the ball rolls when struck. In conventional putter design, such textures are typically achieved through milling the face or using specialized secondary inserts. However, AM can enable complex patterns to be integrated directly into the face itself during the printing process.O-DfAM 4: Multi-material distribution. Related to the use of hierarchical complexity for complex face patterns, multi-material AM can be used to strategically tailor the properties along the face to achieve the desired level of softness or hardness upon impact with the ball. Beyond only the stiffness, multi-material AM can also be used to improve the durability of the face, such as with carbon fiber AM.

Beyond only the stiffness, multi-material AM can also be used to improve the durability of the face, such as with carbon fiber AM. Materials with different densities may also be used to further distribute weight according to the golfer’s preferences.O-DfAM 5: Customization and one-off designs: The putter is the most frequently used club in a round of golf since it is often used on every hole. Because of this, a cottage industry has grown up around the customization of putters, with the intent of allowing golfers an opportunity for personal expression through the design of their clubs. Such customization can be purely aesthetic, such as adding a name or logo, or functional, such as adding customized weighting or patterns to the face of the putter to achieve a specific feel. Customization, as well as other opportunistic DfAM considerations are shown in Figure 3 as they have previously manifested in published AM putter designs.Figure 3.

Customization, as well as other opportunistic DfAM considerations are shown in Figure 3 as they have previously manifested in published AM putter designs.Figure 3. Examples of Opportunistic DfAM in AM Putters (Adapted from [48] and [49])R-DfAM 1: Printer accuracy and tolerances. Transitioning to R-DfAM, the accuracy of the putter’s manufacturing process becomes crucial, assuming the task is to manufacture solely the putter head itself. This is due to the need for the putter head to attach successfully to a provided third-party golf shaft. Though an adhesive is often used to secure a final fit to the shaft, tight tolerances are still needed on the hosel of the putter head. This is especially important in an educational context, where putter heads are swapped in and out during the testing process.R-DfAM 2: Consideration of support material. As with most AM products, support material is a constant consideration for the design and manufacture of a golf putter.

As with most AM products, support material is a constant consideration for the design and manufacture of a golf putter. Barring the simplest designs, it is highly unlikely that any putter generated by students will completely avoid the need2142for support material. This becomes especially relevant as students begin to leverage the shape and hierarchical complexity relevant to opportunistic DfAM in their putter design. Even a traditional blade-style design will require consideration of support material, given the shallow angles present under the heel and toe of the putter head.R-DfAM 3: Limitations on minimum feature size. As with the consideration of support material, all but the simplest putter geometries will be required to account for the minimum feature size achievable by the chosen printing system. For positive features, this may manifest as lattice structures or organic topologies on the main body of the putter head.

For positive features, this may manifest as lattice structures or organic topologies on the main body of the putter head. For negative features, minimum feature size will likely become an important consideration when creating any patterns on the face of the putter to affect the feel of a ball strike. Given the maximum dimensions dictated by the USGA design rules, features in most students’ putter designs will in some way approach the minimum feature size of a printer.R-DfAM 4: Accounting for Anisotropy. As the putter is intended to be an end-use product which will come under load during use (i.e., when striking the ball), students will naturally be required to consider any effects due to material anisotropy that could cause premature failure in the product. In most cases, this will likely be of largest concern in the hosel portion of the putter, which tends to be long and slender and the point at which the pre-built shaft will attach to the printed putter head.

Given other considerations for build time and support material avoidance, the hosel will also likely be built upright, placing the printed layers perpendicular to the direction of the impact force when the ball is struck. This orientation will naturally cause decreased strength that must be accounted for in the design.Figure 4. Examples of Restrictive DfAM in AM Putters (Adapted from [50] and [51])3.3. Demonstrability: End-Use Product That Can Be Created and Tested in the Classroom The third and final characteristic of a successful educational DfAM challenge is the notion of demonstrability. This ultimately drives the ability to not only manufacture student designs easily in the classroom, but also to test them as true-end use products without the need for extensive testing equipment.Build inherently suited for desktop AM size.

While some universities allow students access to large, industrial AM systems, most coursework or individual projects are conducted on small, desktop AM systems, owing to their relative affordability and ease of use. This places a manufacturing size constraint on any design challenge given to student participants. In the case of a golf putter, USGA rules place the largest allowable dimension at 7 inches, which falls2143approximately in line with most desktop-scale AM systems, creating a natural synergy between the required product size and the size of the available manufacturing equipment.Suitable throughput for typical course size. Due to size constraints dictated by the USGA, a typical putter can be easily manufactured overnight or within one 8-hour operation period. This has the potential to enable a generally rapid turnaround time for a single team to design, print, and test their design.

This has the potential to enable a generally rapid turnaround time for a single team to design, print, and test their design. Given an available makerspace or openly available 3D printing facility, this should ensure that student teams are able to generate multiple design iterations on a timetable that works for either a shorter educational module or a semester-long project. However, it is worth noting that throughput could still prove challenging if only one printing system is available or if course enrollment is atypically large.Integrates simply with external components. For the final product to be fully realized for testing, all that is needed is a single external component, that being a standard golf shaft with a grip. Though commercial putters use adhesive to connect the putter head to the shaft, making for a semi-permanent attachment, this is not necessary for testing in a classroom environment.

Though commercial putters use adhesive to connect the putter head to the shaft, making for a semi-permanent attachment, this is not necessary for testing in a classroom environment. Instead, this attachment can be easily mimicked using friction tape affixed to the lower end of the shaft where it inserts into the hosel. In this way, the putter heads can be easily removed and attached to the shaft for testing. Similarly, should students desire to incorporate extra weight into their putter head, via in-situ embedding, a range of simple, everyday items can be used, such as washers, fishing weights, or coins.Easily, quickly testable in a classroom environment. After printing and attaching the putter head to the shaft, the putters can be easily tested in a classroom environment. This can be either on the bare floor, carpet, or, if desired, using a putting mat. Alternatively, for universities with golf courses, students can perform in-situ testing at a practice facility.

This can be either on the bare floor, carpet, or, if desired, using a putting mat. Alternatively, for universities with golf courses, students can perform in-situ testing at a practice facility. The flexibility of this testing enables students to quickly get a feel for the performance and stability of their design, while also gaining feedback from their peers. Given additional space and putter shafts, multiple putters can also be tested simultaneously. Finally, the nature of printing a significant component for an athletic activity naturally introduces the potential to include a competitive element to testing.Viable as end-use product, rather than solely as prototype. It is important that, through DfAM coursework, students begin to understand the validity of the technology as a means for end-use, functional products. This is especially important when students primarily have access to desktop polymer systems, which they may associate with prototypes or aesthetic pieces only.

This is especially important when students primarily have access to desktop polymer systems, which they may associate with prototypes or aesthetic pieces only. For a golf putter, there are no limitations on what the USGA allows for them to be built from. While putters are often made of metal, they can just as viably be made from polymer or wood, due to the relatively low loading they undergo when striking a ball. The result is a manufactured product that may avoid the feeling of students simply working with a “toy” problem or arbitrary academic exercise.4. Implementation in Undergraduate CourseworkTo demonstrate the use of the golf putter design challenge in practice, the remainder of this paper focuses on discussing the implementation of the activity in an undergraduate, DfAM-centered course during the Fall of 2022. The challenge took place over 16 weeks in a course with 242144students at a large, northeastern R1 university.

The challenge took place over 16 weeks in a course with 242144students at a large, northeastern R1 university. Broadly speaking, student participants were tasked to design, print, and test a golf putter head that accounts for restrictive and opportunistic DfAM, where the head must successfully mate to a provided golf shaft and adhere to all USGA requirements. The design challenge in full can be found in the Appendix of this paper. The challenge was introduced at the beginning of the semester, and all topics in Table 1 were covered over the course of the semester, with restrictive DfAM accounting for approximately the first 6 weeks of content and opportunistic DfAM accounting for the remaining 10.To help guide the execution of this project, students were given instructions to perform two iterations of the design. The first iteration was to focus solely on restrictive DfAM, leveraging existing golf putter designs and adapting them to be printable via desktop AM.

The first iteration was to focus solely on restrictive DfAM, leveraging existing golf putter designs and adapting them to be printable via desktop AM. During this iteration, student teams explored the design rules set out by the USGA, the governing organization that drives golf club design. They then conducted an extensive search of existing golf putter designs, including blade, mallet, and half-mallet designs to identify viable options to inform their own designs. Finally, most teams chose what they felt to be the most opportune design and began to adapt it for the characteristics of restrictive DfAM discussed to that point in the semester. At the end of the first iteration, all student teams presented their manufactured designs for review by the rest of the class. Figure 5 shows two such iterations made from black PLA filament.Figure 5.

Figure 5 shows two such iterations made from black PLA filament.Figure 5. Restrictive (Black) and Opportunistic (White) Iterations of Student DesignsThrough this first iteration, student teams came to a shared understanding of how AM limits the design and creation of golf putters. Foremost, student teams realized that PLA, by default, is too light and weak to leverage as a functional putter, even at a 100% infill. The teams that aimed to maximize the mass of PLA in the original design ended the first iteration with a large block of material as a final design. However, when testing, they noted that such large blocks of material do not feel stable or smooth during a swing. Several teams also suffered fracture in their hosel designs due to the anisotropy of the printed material. Additionally, few of the student teams were successful in designing a putter head that was able to fit and secure to the provided golf shaft.

Additionally, few of the student teams were successful in designing a putter head that was able to fit and secure to the provided golf shaft. This reinforced the importance of tolerances when designing for AM.For the second iteration of their designs, student teams were instructed to not only continue to consider restrictive DfAM, accounting for the feedback gained from testing the first iteration, but they were also instructed to introduce the use of opportunistic DfAM to their design. This approach mimics previous research that notes restrictive DfAM should be considered first, followed by opportunistic DfAM to maximize the technical quality of generated design solutions for AM [52]. Beneficially, the limitations students identified in their first iterations synergized well with several opportunistic DfAM characteristics. The first opportunity identified by almost2145every team was that of in-situ embedding for the putter heads.

The first opportunity identified by almost2145every team was that of in-situ embedding for the putter heads. To increase the weight of the initial PLA designs, teams embedded fishing weights, coins, etc. to raise the weight of the putters closer to conventional metal designs. Cavities for such weights can be seen in the parts made from white filament in Figure 5.In conjunction with this, students in several teams investigated the use of geometric complexity to strategically redistribute material with the aim of adjusting the moment of inertia to increase putter stability. Given the lightweight nature of PLA, this often coincided with in-situ embedding; embedded weights were moved to the perimeter of the head where possible. Similarly, during the initial design review, one team experimented with translating a milling texture from a known putter face to test the resolution of their chosen printer.

Similarly, during the initial design review, one team experimented with translating a milling texture from a known putter face to test the resolution of their chosen printer. After this review, the remaining teams made intentional efforts to likewise include hierarchical complexity as a means of adjusting the texture of the putter face to improve putting feel. Further, one team combined this complexity with the use of modular customization into their putter through a series of removable putter faces with custom designed hierarchical textures (seen in the leftmost white printed part in Figure 5). Finally, one team customized/personalized the bottom of the putter with an imprint of a university mascot; this team’s design (seen in Figure 6), also incorporated geometric complexity, in-situ embedding, and hierarchical complexity, alongside a series of restrictive DfAM considerations.

This helps to demonstrate the ultimate potential of this design challenge in practice; where students can organically incorporate a range of opportunistic and restrictive DfAM considerations into a true end-use product. Note that, while no teams leveraged multi-material printing, several expressed interest. However, the one accessible multi- material extrusion printer was out of service at the time of the second iteration.Figure 6. Restrictive (Grey) and Opportunistic (Blue) Iterations of Winning Student Design5. Conclusions and Future WorkIn this paper, an argument has been made for using a golf putter design challenge as an educational catalyst for DfAM via PBL. This challenge offers clear design guidelines, a simple to understand end-use context, a robust need for both restrictive and opportunistic DfAM concepts, and is easily manufactured and tested in a classroom environment.

This leads to one of the fundamental contributions of this paper: the notion that, to be a successful design challenge for DfAM PBL, a task must adhere to the three principles of clarity, applicability, and demonstrability. If a design challenge falls short in one of these three aspects, then there is the2146possibility that the DfAM instructor or students may encounter unnecessary roadblocks in the implementation and execution of the challenge, hindering its ultimate usefulness as an educational tool. To demonstrate the usefulness of the golf putter challenge in practice, its implementation in an upper-level undergraduate design course was shown, along with example student designs and lessons learned throughout the progression of the challenge over the course of the semester.While initial efforts at deploying the golf putter were promising, this research is still a work in progress.

Future work requires more formal and rigorous evaluation of the deployed design challenge and its effect on student design outcomes and learning. Further, it is necessary to understand how variations in the putter challenge (e.g., explicit constraints/objectives, team- based vs. individual project, etc.) and student demographics (e.g., year of study, existing knowledge of DfAM) may change its effectiveness. Lastly, the design challenge must be directly compared against other existing design challenges to validate its usefulness against the wide range of tasks already available in support of DfAM PBL.6. References[1][2]Dym, C. L., Agogino, A. M., Eris, O., Frey, D. D., and Leifer, L. J., 2005, “Engineering Design Thinking,” J. Engingeering Educ., (January), pp. 103–120. Hendry, G. D., Frommer, M., and Walker, R. A., 1999, “Constructivism and Problem‐based Learning,” J. Furth. High. Educ., 23(3), pp. 369–371.[3] Williams, C. B., and Seepersad, C.

[Accessed: 24-May- 2023].[49] Sherman, J., 2018, “Is 3D Printing the Future of Putter Design? Kronos Golf ThinksSo...,” Pract. Golf [Online]. Available: https://practical-golf.com/kronos-golf-3d-printing/. [Accessed: 24-May-2023].[50] GE Additive, 2018, “GE Additive Helps Callaway Golf Company to Redesign PutterHead” [Online]. Available: https://www.ge.com/additive/press-releases/ge-additive-helps- callaway-golf-company-redesign-putter-head. [Accessed: 24-May-2023].[51] Thimmesch, D., 2015, “Golf Equipment Manufacturer Ping Introduces Golf’s First 3D- Printed Putter,” 3Dprint.com [Online]. Available: https://3dprint.com/46036/golf- equipment-manufacturer-ping-introduces-golfs-first-3d-printed-putter/. [Accessed: 24- May-2023].[52] Prabhu, R., Simpson, T. W., Miller, S. R., and Meisel, N. A., 2021, “Fresh in My Mind! Investigating the Effects of the Order of Presenting Opportunistic and Restrictive Design for Additive Manufacturing Content on Students’ Creativity,” J. Eng.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesign guidelines for a software-supported adaptation of additively manufactured components with regard to a robust productionS. Lammers1,2, J. Tominski1,2, S. Magerkohl1,2, T. Künneke1,2, T. Lieneke1,2, D. Zimmer1,21Chair of Design and Drive Technology, Paderborn University, 33098 Paderborn, GermanyDirect Manufacturing Research Center (DMRC), Paderborn University, 33098 Paderborn, GermanyAbstractThe design of additively manufactured components requires a rethinking in the design process. This is inhibited by a lack of knowledge about additive manufacturing technologies. For this reason, a large number of design guidelines have been developed in recent years.

This is inhibited by a lack of knowledge about additive manufacturing technologies. For this reason, a large number of design guidelines have been developed in recent years. In their present form the design guidelines are not suitable for processing in a software algorithm, since the guidelines have a certain redundancy and partly influence each other. This paper describes several steps to consolidate the existing guidelines and to prepare them in a way that they can be used in a software algorithm for a design check. Therefore, existing guidelines are collected, prioritized and quantified with regard to their relevance for a robust production. To quantify the guidelines, test specimens are developed, produced and evaluated in order to obtain a limit value for the geometric properties.

To quantify the guidelines, test specimens are developed, produced and evaluated in order to obtain a limit value for the geometric properties. With these limit values, quantifiable design guidelines can be applied to designers and software tools.IntroductionAdditive manufacturing (AM) processes are increasingly used in the production of components and structures in industrial practice. The term additive manufacturing refers to processes in which a component is manufactured in layers and without forming tools [Tho09, Geb07, VDI3405, LQM+17, LAS+16]. The potentials of the layered design is visible from an economic and technical point of view. In particular, the freedom in the geometric design of components enables to produce structures that cannot be produced using conventional manufacturing processes such as turning or milling. Complex structures, such as lattice structures or bionic and topologically optimized structures can be produced.

Complex structures, such as lattice structures or bionic and topologically optimized structures can be produced. However, additive manufacturing technology is also subject to some restrictions. The challenge of this relatively new technology is to exploit the potential mentioned, taking into account the process-specific restrictions. New design processes and possibilities in component design require a rethinking, especially in the design process. This rethinking is made more difficult by a lack of knowledge of additive manufacturing technology [HHD06, WC15].In order to make efficient use of AM, design recommendations have been made in recent years [Ada15, KHE15, Tho09, VDI3405-3]. These design recommendations support the design for production and thereby the minimizing of geometric deviations during the design process in order to enable a robust manufacturing.

These design recommendations support the design for production and thereby the minimizing of geometric deviations during the design process in order to enable a robust manufacturing. Since the development has resulted in a large number of recommendations, which contain different quantitative information, it is increasingly difficult to maintain an overview and to identify the most important recommendations concerning a robust production.Therefore the present study focuses on an automated, software-based design check for additively manufactured components with laser beam melting (LBM) with regard to manufacturability, minimal post-processing by removing support material and the best possible compliance with geometric requirements. An automated, software-based design check should support the conception and design of AM components. Furthermore, it is possible to test lattice structures and topologically optimized components for a robust production.

Furthermore, it is possible to test lattice structures and topologically optimized components for a robust production. For this reason, existing design recommendations have to be reviewed and prioritized according to their527relevance for the AM process and component quality. The aim is to identify a reduced number of recommendations that include all process-specific restrictions. These recommendations should then be prepared and experimentally tested in order to obtain relevant design guidelines for the respective case, which can be used in an automated software-based design check.State of the art: Design recommendationsThe scientific literature contains numerous papers containing recommendations for the design of AM components suitable for production.

Some work dealing with the laser beam melting process will be presented below.In his work on various component geometries, THOMAS [Tho09] carries out experimental parameter studies with the material 316L and analyses these with regard to production-related restrictions. The results of the analysis form the basis for the preparation of design recommendations. For this purpose, sample geometries are developed to specifically evoke previously observed effects. Through targeted variation of geometric dimensions, THOMAS determines how a geometric property is to be designed in order to produce it reliably using laser beam melting processes. His recommendations refer in particular to the design of overhangs, gaps, bores and the need for support structures.ADAM [Ada15] develops design recommendations for fused deposition modeling, laser sintering and laser beam melting. He considers so-called geometric standard elements, which are used repeatedly in component design.

He considers so-called geometric standard elements, which are used repeatedly in component design. Based on these standard elements, he develops specimen geometries that reflect the properties of these standard elements. By varying the geometric properties, he identifies conditions in which robust production can be guaranteed no longer. ADAM records the findings of its investigation in recommendations that describe a fact quantitatively or qualitatively.[KHE15] carry out experimental investigations with the material TiAl6V4. The aim of the work is to define design recommendations for lightweight components so that they can be manufactured and used reliably. Typical elements that are frequently recurring in lightweight design are considered, such as beam elements and cavities. The results are classified in tabular form and supplemented by qualitative and quantitative data.KRANZ, HERZOG and EMMELMANNSome of the literature contains recommendations describing the same facts.

However, the quantitative data often differ due to different materials, machines and their parameter settings. Furthermore, some of the work contains recommendations whose essence can also be applied to other recommendations and geometries.For this reason, existing design guidelines need to be consolidated and prioritized to reduce their number for a software-based design check. Furthermore, they need to be mathematically describable so that they can be processed in a software. This poses a particular challenge for design recommendations with qualitative data, which have to be transformed into quantitative data.Priorization of design recommendationsThe knowledge about a production-ready design of AM components is manifold and influenced by many factors. Most publications on design recommendations for AM processes are often comparatively abstract and therefore less qualified for the generation of quantitative indications.

Most publications on design recommendations for AM processes are often comparatively abstract and therefore less qualified for the generation of quantitative indications. In order to make these recommendations manageable for a software based design check, a procedure is developed, which provides a compact overview of relevant design recommendations and quantifies characteristic values. In the course of this, the design recommendations mentioned above [Ada15, KHE15, Tho09] are arranged in a tabular overview. Despite diverging boundary conditions in the derivation of design recommendations,528agreement can be found among the different authors. By prioritizing these nearly 100 design recommendations, relevant guidelines for robust production can be identified.

By prioritizing these nearly 100 design recommendations, relevant guidelines for robust production can be identified. The design recommendations are allocated to four priority levels according to Table 1.Table 1:Prioritization scheme of the design recommendationsPriority level 1DescriptionViolation of the design recommendation endangers the manufacturability or functionality of the component.Exemplary design recommendationsAdherence to the minimum wall thickness to ensure the design of a wall. [KHE15]2Violation of the design recommendation deteriorates the component properties (surface quality, geometric properties).

[KHE15]2Violation of the design recommendation deteriorates the component properties (surface quality, geometric properties). These deteriorations can be corrected by post-processing.Provision of surfaces oriented perpendicular to the building platform in order to avoid the stair step effect, as this worsens the component surface [Ada15].3Design recommendations enable users to determine the dimensions or positions of the components themselves without setting a specified limit value. They have no significant influence on component quality.The user can freely select the component position in the building chamber. [Ada15]4Design recommendations are also taken into account in other design recommendations.Overhanging sections with an angle of <50° must be supported [KHE15]. This recommendation is already included in the recommendation to orientate surfaces 45° to the building platform.

This recommendation is already included in the recommendation to orientate surfaces 45° to the building platform. [Ada15]The recommendations classified in priority level 1 and 2 are most important for use in a software for production-oriented design checking and have a direct influence on a robust production, component functionality and component quality. Due to their relevance for automated, software-based design verification, these recommendations will be called design guidelines in the following. These are described in Tabels 2-9 where they are categorized in eight superior categories. They are prepared graphically by listing the essential effects. These are shown in the second and the third column, respectively, by an example suitable for production and an example not suitable for production. Each guideline is also listed in written form in the first column.

Backgrounds are explained in the surrounding text.Surface orientationThe guideline category for surface orientation contains four effects, that occur when theorientation angle δO,lim falls below a limit value:• a production-related oversize is introduced by the surfaces orientated obliquely to the building platform, because of powder adheres due to the laser penetration depth beyond a component layer 529the surface quality is reduced by the need for support.the stair step effect, which results from the layered construction and the approximation of the component contour through rectangular layers which also leads to a poorer surface quality at low surface orientation angles.These effects are mapped to priority level 2.a robust production is endangered if the surface orientation angle of unsupported surfaces falls below a limit value.This effect has to be allocated to priority level 1Table 2:Guideline “Surface orientiation“Not suitable for productionSuitable for production1 Surface orientation 1.1 Support: Angledsurfaces must not fall below the angle δO,lim relative to the building platform to avoid support1.2 Powder adherence: Unintentional powder buildup with a thickness aO occurs on the surfaces facing downwards, which cannot be prevented.

The dimension aO must be taken into account in the design so that the actual dimension ai complies with the nominal dimension an. 1.3 Stair step effect: If possible, large surfaces should be aligned orthogonally to the building plane (δO = 90°).1.4 Unsupported surfaces: Unsupported surfaces should not fall below a limit value for δO to ensure a robust productionOuter radiusThe minimum possible outer radius is limited depending on the orientation in the building chamber. Thus, the radius in 90° to the building platform is dependent on the formation530of a contour path and an inner grid, which only occurs at a limit value for the radius ro,90,lim. The cross-section is also limited in 0° orientation and should not fall below a limit value to allow the component to detach from the building platform. These guidelines are allocated to priority level 1. Furthermore, the stair step effect occurs in rounded component areas at a limit value for the radius.

These guidelines are allocated to priority level 1. Furthermore, the stair step effect occurs in rounded component areas at a limit value for the radius. These guideline is mapped to level 2.Table 3:Guideline ”Outer radius“Not suitable for productionSuitable for production2 Outer radius 2.1 Minimum 90°: The outer radius ro,90 of cylindrical components must be greater than or equal to the permissible outer radius ro,90,lim.2.2 Minimum 0°: The outer radius ro,0 of cylindrical elements must be greater than or equal to the permissible outer radius ro,0,lim.2.3 Stair step effect: The outer radius ro,sse should exceed the permissible outer radius ro,sse,lim.Inner radiusThe inner radius cannot be formed with very small radii, as powder adhesion results in radii that are below the limit value ri,90,lim. This effect occurs both in the 90° and in the 0° orientation. Moreover unsupported surfaces can only be build until a limit radius.

This effect occurs both in the 90° and in the 0° orientation. Moreover unsupported surfaces can only be build until a limit radius. These three effects are allocated to level 1. In addition, rounded surfaces are approximated by rectangular surfaces, which lead to a deterioration of the surface quality due to the stair step effect. The deterioration of the surface can be increased by the need for support above a limit value for the radius.

The deterioration of the surface can be increased by the need for support above a limit value for the radius. Because of its relevance for robust production, this guideline is are mapped to priority level 1.531Table 4:Guideline “Inner radius“Not suitable for productionSuitable for production3 Inner radius 3.1 Minimum 90°: Cylindrical cavities oriented perpendicular to the building platform need an inner radius of ri,90,lim.3.2 Minimum 0°: Cylindrical cavities, which are not oriented perpendicular to the building platform need an inner radius of ri,0,lim.3.3 Stair step effect: The inner radius ri,sse should be larger than the permissible inner radius ri,sse,lim.3.4 Support: To produce a cavity or an overhang with the inner radius ri,s without support structures, the radius ri,s of the downward facing surface must be selected smaller than or equal to the permissible inner radius ri,s,lim.

3.5 Unsupported Surfaces: Unsupported surfaces should not increase over a limit value to ensure a robust production without supportCross-sectional area change in building directionWhile building components in layers, it must be ensured that the ratio of the maximum cross-sectional area Amax and the minimum cross-sectional area Amin is less than or equal to the532factor fA, especially for branched components. This applies in the case that Amax is in a higher layer than Amin. When determining the cross-sectional areas Amin and Amax, not only the cross- sectional area of the solid component but also the cross-sectional areas of solid support structures are added up.

When determining the cross-sectional areas Amin and Amax, not only the cross- sectional area of the solid component but also the cross-sectional areas of solid support structures are added up. This guideline is to be assigned to priority level 1, as non-compliance will lead to a endangering of the build job.Table 5:Guideline ”Cross-sectional area change in building direction“Not suitable for productionSuitable for production4 Cross-sectional area change in building direction The ratio of the cross- sectional area of Amax to Amin should not exceed the factor fA over the overall height z.Wall thicknessThese guidelines are mapped to priority level 1, as violation leads to walls not being built up either in 0° or 90° orientation. This is on the one hand due to the lack of contour and grid paths in the 90° orientation.

This is on the one hand due to the lack of contour and grid paths in the 90° orientation. And on the other hand to the fact that walls in the 0°orientation under a limit wall thickness cannot be detached from the support and have holes in the surface.Table 6:Guideline “Wall thickness“Not suitable for productionSuitable for production5 Wall thickness 5.1 Minimum 90°: The wall thickness tw,90 must not be less than the minimum wall thickness tw,90,lim.5.2 Minimum 0°: The wall thickness tw,0 must not be less than a minimum wall thickness tw,0,lim.533Gap widthGaps are not formed below a minimal gap width, as the gap is closed by powderadhesion. Violation of the rule therefore leads to a loss of function.

Violation of the rule therefore leads to a loss of function. For this reason, the rule must be assigned to priority level 1.Table 7:Guideline “Gap width“Not suitable for productionSuitable for production6 Gap width 6.1 Minimum 90°: Gaps arranged perpendicular to the building platform or x/y-plane must have at least a gap width wg,90 of wg,90,lim.6.2 Minimum 45°: Gaps arranged at an angle to the building platform must have at least a gap width wg,45 of wg,45,limOverhangLong overhangs increase the area where heat penetrates the component without creating additional space for heat transport to the building platform. This can lead to heat accumulation, which leads to shrinking processes in the component. The resulting internal stresses can lead to plastic deformation of the component and termination of the build job.

The resulting internal stresses can lead to plastic deformation of the component and termination of the build job. For this reason, the guideline has to be mapped to priority level 1.Table 8: Guideline “Overhang“Not suitable for productionSuitable for production7 Overhang 7.1 Support: The length of an overhang in the xy- plane must not exceed l0,lim without the use of support material.534Cross-section areaCross-sectional areas in a component layer should not exceed a limit value, since the layerwise melting of powders results in locally different temperature gradients. This can lead to residual stresses and therefore to deformations and endanger the build job.

This can lead to residual stresses and therefore to deformations and endanger the build job. For this reason, these guidelines are allocated to priority level 1.Table 9: Guideline ”Cross section area“Not suitable for productionSuitable for production8 Cross-section area 8.1 Maximum (cubic): The continuous rectangular cross- sectional area Ac in the x-y-plane must not exceed the critical area Ac,lim.8.2 Maximum (round): The continuous round cross-sectional area Ar in the x-y-plane must not exceed the critical area Ar,lim.Quantification of design rulesTo quantify the guidelines, test specimens were developed for each guideline category. The test specimens allow the verification of the respective attributes, which are considered in the corresponding guidelines. For this purpose, the test specimens were produced under different machine-, material-, and production parameter combinations.Further on the permissible limit value for the attribute was determined.

The test specimens were divided into two specimen build jobs, which are referred as test specimen build job 1 and 2. Test specimen build job 1 contains test specimens that have to be sorted as critical with regard to a robust production. These test specimens can lead to an abort of the build job. A successful or unsuccessful production of the test specimens can be used as an evaluation criterion. Specimen build job 2 contains specimens that can be manufactured robustly. These specimens can be measured optically or tactilely after production [LTM+18, TL18].The test specimens were manufactured for all guidelines with 316L stainless steel (1.4404) and AlSi10Mg aluminium alloy on a SLM 250 HL and SLM 280 HL systems. For each machine-material combination, the test specimens were produced for two common parameter sets. A total of eight build jobs with different boundary condition combinations (machine, material, parameter set) were produced.

A total of eight build jobs with different boundary condition combinations (machine, material, parameter set) were produced. These eight build jobs were each produced in triplicate, resulting in a total number of 24 build jobs. The parameter sets (PS) used, are listed in Table 10.535Table 10:Parameter set for the production of the specimens.Machine SLM 250 HLSLM 280 HLMaterial 316LAlSi10Mg316LAlSi10MgParameter set PS1: 316L_S09-02A16-312V3_30 PS2: 316L_S32-14A18-506V1_50 PS1: AlSi10Mg_S32-14A20-514V4_30 PS2: AlSi10Mg_S32-14A18-512V5_50 PS1: 316L_SLM_BP2.1_V5104_30 PS2: 316L_SLM_BP2.1_V5104_50 PS1: Al_SLM_BP2.1_V5102_30 PS2: Al_SLM_BP2.1_V5102_30The buildjob preparation was carried out with Materialise Magics V21.1. In case of the 250 HL Autofab 1.8 was used for the slicing process. After production, the test specimens were separated from the building platform with a band saw.

In case of the 250 HL Autofab 1.8 was used for the slicing process. After production, the test specimens were separated from the building platform with a band saw. The subsequent measurement is explained exemplary for the test specimens for guidelines 5 (wall thickness) and 6 (gap width). The corresponding test specimens and determined limit values for different boundary condition combinations are listed in the tables 11 and 12. The limit values and test specimens of the remaining design guidelines are listed in the appendix.For the test specimen of guideline 5 the minimum wall thickness shall be determined from which a robust manufacturing of the wall is possible. The wall thickness tw,90 and tw,0 is varied in the range of 0.1mm to 1.0mm in 0.1mm steps. The wall surfaces are then subjected to a visual inspection and evaluated with a view to robust production. Surfaces that are perforated or walls that have not been formed out are declared as not possible to manufacture.

Surfaces that are perforated or walls that have not been formed out are declared as not possible to manufacture. The maximum wall thickness which enables such a declaration must at least be maintained in order to be able to manufacture a wall.In addition, the target wall thickness tw,90 was measured at the upper two corner points (CP1, CP2) of the vertical walls using an outside micrometer to detect geometric deviations. The measured values on the triple specimens are recorded in tabular form and averaged for the different boundary conditions to determine the corresponding limit values. The limit values are shown in Table 11. Both machines used for the production show the same limit values for 316L. AlSi10 Mg diverge from this results and show a higher limit value. This is due to the greater thermal conductivity of the aluminium and the associated greater heat input into the powder bed, which leads to the melting of powder particles onto the component contour.

This is due to the greater thermal conductivity of the aluminium and the associated greater heat input into the powder bed, which leads to the melting of powder particles onto the component contour. Besides this is caused by bigger adhesions related to different powder particle size of AlSi10Mg.Table 11:Specimen and limit values for tw in design guideline 5 “Wall thickness”MachineParameter -setMaterial316LAlSi10MgSLM 250PS1≥0,4 mm≥0,3 mmPS2≥0,4 mm≥0,4 mmSLM 280PS1≥0,4 mm≥0,4 mmPS2≥0,4 mm≥0,3 mm536The test specimens for guideline 6 shall be used to determine the permissible gap width perpendicular to the building platform (wg,90) and at an angle of 45° (wg,45) to the building platform (see Table 12). The gap width was varied in steps of 0.05mm from 0.05mm to 0.3 mm. In this case, the measurement is carried out with a feeler gauge. It is checked whether the complete gap is free by pushing the largest possible feeler gauge into the gap.

In this case, the measurement is carried out with a feeler gauge. It is checked whether the complete gap is free by pushing the largest possible feeler gauge into the gap. The feeler gauge must be inserted completely into the gap. The maximum possible insertable thickness of the feeler gauge is recorded as the permissible limit value. The determined limit values are listed in Table 12. It is recognizable, that the SLM 280 parameters allow a smaller minimum gap width of 0.15mm for both materials. This is due to the different process parameters of the machines.Table 12:Specimen and limit values for wg in design guideline 6 “Gap width”.MachineParameter -setMaterial316LAlSi10MgSLM 250PS1≥0,15 mm≥0,2 mmPS2≥0,3 mm≥0,3 mmSLM 280PS1≥0,15 mm≥0,3 mmPS2≥0,3 mm≥0,3 mmSummary / OutlookIn order to be able to integrate the existing design recommendations into a software, a systematic review and prioritization of these is required.

Approximately 100 design guidelines were examined. These were evaluated with a view to robust production. Overlapping and similar design guidelines were summarized. This reduced the number of design guidelines most important for a software-based design check to eight. These guidelines are described in detail and discussed using sample graphical illustrations.Test specimens were developed to determine permissible quantifiable limit values for the attributes of the design guideline. For this purpose, the test specimens were manufactured and examined as well as the permissible limit values determined. These limits values represent the basis for a software-based design check. These were produced for different boundary conditions (machine, materials, parameter sets) to enable applicability.

These limits values represent the basis for a software-based design check. These were produced for different boundary conditions (machine, materials, parameter sets) to enable applicability. The procedure for determining limit values was described in detail as well to enable the user to determine limit values for his own boundary conditions.In the future, limit values for further boundary condition combinations should be determined in order to enable a broad application of the design guidelines. In addition, these limits are to be used for software-supported design testing and implemented in a software tool. Furthermore, guidelines for a post-processing, load-fair and cost-fair design are to be determined and implemented in a design checking software.The research leading to these results has received funding from the BMBF projectOptiAMix – “Multi-target-optimized and continuously automated product development for additive manufacturing in the product development process”.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceA Design Method to Exploit Synergies Between Fiber-Reinforce Composites and Additive Manufactured ProcessesVahid Hassani1, David W. Rosen1,2, Ethan Goh3, Sufiyan Sarwan3, Florian Doetzer31Digital Manufacturing and Design Centre, Singapore University of Technology and Design 2George W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332 3Composite Cluster SingaporeCorresponding author. Email: david.rosen@me.gatech.eduAbstractThis paper proposes a design method for devices composed of long fiber-reinforced composites (FRC) and additive manufactured (AM) parts. Both FRC and AM processes have similar application characteristics: suitable for small production volumes, additive in nature, and capable of being highly automated. On the other hand, the classes have distinct characteristics.

On the other hand, the classes have distinct characteristics. FRC components tend to be large and of simple shapes, while AM components tend to be small with highly complex geometry. Their combination has the potential for significant synergies, while mitigating their individual limitations. A decision guide is proposed, in the form of a series of questions, to guide the designer to determine if their application is a good candidate for FRC+AM. The decision guide is reformulated into a proposed design process that guides the designer to advantageously benefit from AM and FRC characteristics. The tools are illustrated with an example of a composite pressure vessel with integrated pressure reducer.1 BackgroundThis paper proposes a design method for devices composed of long fiber-reinforced composites (FRC) and additive manufactured (AM) parts.

Both FRP and AM processes have similar application characteristics such as being suitable for small production volumes, being additive in nature, and capable of being highly automated. On the other hand, the two classes of processes have distinct characteristics. FRP components are superior for large structures with simple shapes and load cases [1], where directed fiber optimizations can be exploited fully. However, composites are difficult to use for complex components, because of the inherent design and manufacturing complexity. AM excels for highly complex components and customization, but usually are limited in size [2]. Regarding their manufacturing processes, composite processing is becoming more automated, making the approach somewhat similar to AM [3,4]. Manufacturing lot sizes for composites and AM have traditionally been low, but with technological advances are now reaching up to tens of thousands.

Manufacturing lot sizes for composites and AM have traditionally been low, but with technological advances are now reaching up to tens of thousands. Therefore, the combination of both technologies promises enormous potential, if experience in their integrated design can be codified and disseminated and design guidelines established.The authors have many years of design and fabrication experience with either FRC or AM. They recently completed a significant research project to investigate opportunities to leverage the unique capabilities of FRC and AM, and overcome their limitations. As a result, both a decision guide and a design process were proposed for FRC+AM devices; that is, devices composed of components manufactured by either AM or FRC. The proposed decision guide is a structured series of questions that is intended to guide the designer to determine if their application is a good candidate for FRC+AM.

The proposed decision guide is a structured series of questions that is intended to guide the designer to determine if their application is a good candidate for FRC+AM. The2015decision guide questions are reformulated into a proposed design process that can guide the designer to advantageously benefit from AM and FRC characteristics. The decision guide and design process are illustrated with an example of a composite pressure vessel with integrated pressure reducer.The specific pressure vessel under study, and its construction, is shown in Figure 1. The composite pressure vessel (CPV) is manufactured by filament winding of carbon fiber tapes impregnated with polyamide-6 (PA6) thermoplastic. The CPV construction consists of the pressure vessel body, end-caps at both ends, and an internal liner. The combination of the pressure vessel body, liner, and end-caps serves as the mandrel on which filament winding is performed. These CPVs have pressure ratings up to 350 bar (5000 psi).

These CPVs have pressure ratings up to 350 bar (5000 psi). The design objective investigated in the research project was to determine if part or all of the pressure reducer could be redesigned and integrated into an end-cap, which would be manufactured via AM.Pressurea)Camon Fiber Reinforced TP Tapet;;;~;;;;::'\==\PA6 End Caps Connect To Adjacent Structural ComponentsLineri~~==-=::1/ :rlb)Figure 1: Cylindrical polymer composite pressure vessel a) with standard pressure reducer, and b) showing internal configuration.2 Composites and AM ProcessesAn introduction to polymer composites manufacturing and AM processes will be provided to serveas a baseline of understanding for the DFM presentation.2.1 Composites manufacturing processesComposite structures are superior for large / extended structures and simple load case scenarios, where directed fiber optimizations can be exploited fully.

However, composites are difficult to use for complex components, because of the inherent design & manufacturing complexity. Composite manufacturing is becoming more and more automated, making the approach somewhat similar to AM (composite manufacturing is always additive). Manufacturing lot sizes for composites and AM have traditionally been low (handful to few hundreds), but with technological advances are now reaching up to tens of thousands.A wide variety of long fiber reinforced polymer composite manufacturing processes have been developed [5]. They are all additive in nature in that fibers, tows, or mats are impregnated with polymer2016and deposited onto the workpiece. Typically, a mold or mandrel is used to provide the desired part shape and the fiber-polymer material is deposited onto the tool or onto already-deposited material. Common composites manufacturing processes include [6]:sheet forming: a flat sheet of composite material is formed by pressing in dies.

Common composites manufacturing processes include [6]:sheet forming: a flat sheet of composite material is formed by pressing in dies. This is an analogous process to sheet metal forming or stamping.filament winding: fibers and polymer matrix are wound onto a rotating mandrel to form the part.resin transfer molding: fiber reinforcement is formed in the mold cavity, then thermosetting resin is forced into the mold at high pressure. A variation is to utilize a vacuum to assist resin infusion.pultrusion: fibers or braids are pulled through a resin bath and a die to fabricate constant cross- section components (combination of pull and extrusion)What these processes have in common is the capability of forming a wide variety of part geometries by adding and processing material using hard tooling.

Overall part geometries can be fabricated but, typically, small details for assembly and joining, or for local functional purposes, cannot be fabricated without secondary operations.For thermosetting polymers, an additional step is often needed for the polymerization reaction to occur. Different types of polymer matrix infusion or infiltration are used. However, these aspects of the processes are not important for this paper.2.2 Filament windingSince filament winding will be used in the case study for this paper, additional information about the process will be presented. As mentioned, in this process, a filament consisting of fibers and polymer matrix is deposited onto a mandrel as the mandrel rotates on a spindle. A carriage containing the deposition head translates back-and-forth along the mandrel, resulting in the filament being wound onto the mandrel.

A carriage containing the deposition head translates back-and-forth along the mandrel, resulting in the filament being wound onto the mandrel. Most filament winding machines incorporate only these two degrees-of-freedom: spindle rotation and carriage translation along the spindle. Some high-end filament winding machines incorporate additional degrees-of-freedom in the deposition head to enable more motion options and deposition patterns.Most often, several layers of filament will be deposited to fabricate the part. Filament deposition angle can be varied by varying the ratio of spindle rotation to carriage translation speeds. High angles (close to 90 degrees relative to the spindle axis) yield parts with high circumferential strength, while lower angles result in parts with greater axial strength.

High angles (close to 90 degrees relative to the spindle axis) yield parts with high circumferential strength, while lower angles result in parts with greater axial strength. Many times, deposition angles will be varied in different layers in order to impart good overall strength characteristics [7].Since parts are fabricated on a rotating spindle with two degree-of-freedom machines, parts are typically simple shapes, such as cylinders. Pressure vessels with hemispherical ends can be fabricated. Since the mandrel is incorporated into the spindle, part ends cannot be completely closed by the filament winding process. As a result, end-caps or some other means of closing ends are needed.Figure 2 shows a schematic of the spindle-mandrel assembly and filament winding machine that was used for pressure vessel fabrication for this paper. In the schematic, the part labeled “Liner” and the Endcaps comprise the mandrel.2017~bQ):.,steel fixture~a)~ ~Threodedinserlb)Figure 2.

In the schematic, the part labeled “Liner” and the Endcaps comprise the mandrel.2017~bQ):.,steel fixture~a)~ ~Threodedinserlb)Figure 2. Filament winding process for the composite pressure vessels, showing a) a schematic of the spindle-mandrel assembly and b) pressure vessel fabrication on a filament winding machine.2.3 Additive manufacturing processesThe AM processes of interest in this paper fabricate parts in polymer and polymer composite materials. Of most interest are material extrusion, material jetting, and powder bed fusion processes, although vat photopolymerization and binder jetting processes could be utilized as well. As is well known, AM processes fabricate parts in a layer-by-layer manner by one of two routes: depositing and processing material, or depositing energy into a material vat in order to process the material.

Material extrusion and jetting are examples of the former, while powder bed fusion utilizes the latter.In contrast to polymer composites, AM machines tend to have smaller build volumes resulting in smaller parts. AM processes are capable of fabricating parts with complex geometries. This manifests in the capabilities of:combining several conventionally manufactured parts into one AM part,fabricating cellular structures, such as lattices and foamsfabricating parts with integral assembly and joining featuresfabricating custom partsTo date, the material extrusion process has been adapted to fabricate long fiber reinforced composites. Two machine vendors, Markforged and Arevo, have developed systems that deposit both thermoplastic filament and a fiber from two deposition heads. In the Markforged system, fibers are deposited in the same 2D layers that the polymer filaments are deposited in, while Arevo claims to enable 3D deposition.

In the Markforged system, fibers are deposited in the same 2D layers that the polymer filaments are deposited in, while Arevo claims to enable 3D deposition. Research systems are being developed that utilize automated fiber placement heads on robotic arms that enable 3D deposition that promise large, complex geometry, 3D fiber reinforced composite parts.2018An increasing interest in design for AM (DFAM) is indicated by a variety of proposed DFAM methods in recent years. Principles for DFAM have been proposed [8] and DFAM methods have been developed to take advantage of AM opportunities [9], as well as design rules for designing around manufacturing process constraints [10].3FRP+AM Decision GuideBased on a structured design thinking process and supported by experiences with composites and AM, a decision guide has been developed to support the design of integrated additive manufacturing and composite components and modules.

The objectives of the decision guide are to help the designer identify whether FRP+AM is suitable for their application and, if so, to identify which parts of the application are suitable for composites and which are suitable for AM.The guide is presented here as a series of questions and actions to perform.1. Are polymer or polymer composite materials suitable for the intended application? 2. 3. Look for opportunities to integrate additional modules into the artifact. 4. Does the artifact, or could the artifact, have geometric detail too small for typical composite manufacturing processes?Is the artifact large enough to be a candidate for composites?5. Can the geometric detail be isolated into specific regions, where other regions have simple shapes?6.

9. 10. Can part designs be modified to facilitate composites manufacturing processes; e.g., add sacrificial material for fixturing, add fixturing features, etc.11. Compare part designs with functional, economic, and sourcing requirements4FRC+AM Design MethodBased on the decision guide questions and thought process, a flowchart version of a design process is presented in Figure 3. The proposed design method starts with a device configuration which has been developed already, based on the decision guide just presented. Conceptual and embodiment design stages should be conducted to identify the primary modules with some preliminary attempts performed to relate functionality to specific geometric regions. Following the decision procedure, some geometric regions should be of simple shape with a small number of functions allocated to them, with the idea that these will be candidates for composites manufacturing processes.

Other geometric regions should be smaller in size with a higher concentration of functional elements. These should be candidates for AM.2019The proposed method proceeds to guide the designer to focus on each geometric region individually and to develop detailed designs for each. Materials and manufacturing processes should be identified and design for manufacturing performed. If multiple parts are needed in a region, each should be designed to a similar level of detail. As necessary, the designer should ensure the soundness of their functional module designs if they consist of more than one geometric region.After the design of all modules, the designer should assemble the modules into a complete system design. The right-most sequence of activities, starting at marker B, should be performed. The designer should investigate the assembly of components and identify suitable joining methods.

The right-most sequence of activities, starting at marker B, should be performed. The designer should investigate the assembly of components and identify suitable joining methods. Concurrently, the designer should consider compatibility among materials and determine if material substitutions should be made for either the composites or AM parts. When joining and assembly methods have been identified, part designs should be refined to incorporate any additional assembly or joining features that are needed to facilitate assembly, such as chamfers (for insertion), snap fits, steps or notches to ensure parts are aligned correctly, etc.As appropriate, the designer should consider costs, delivery times, and other aspects of sourcing for each component and for the entire system.

These considerations should be compared with requirements or targets and the design process iterated if needed.Artifact with geometric regions identifiedFor each geometric region :Refine selection of materials to ensure compatibilityIdentify j oining methodsDesign part for composites processesDesign part or module for AM processesIntegrate assembly and joining features into part designsIdentifymatrix and fiber materials composites man ufacturing processesIdentifyAM materials - AM processesAdd featu res to facilitate composites manufacturingAAssemble region designs into system configurationCompare part designs with functional, economic, and sourcing requirementsConsider other manufacturing process esDevelop economic and sourcing requirementsFigure 3.

Flowchart of proposed composites+AM design process.20205 Composite Pressure VesselIn this section, the example will be present of the composites+AM design process, beginning with a description of the application. The decision procedure presented in Section 3 will be applied to this design problem, which identifies the opportunity to integrate the pressure reducer into an end-cap of the composite pressure vessel (CPV).

The design of the integrated end-cap is used to illustrate the application of the proposed design process.5.1 Description of ApplicationThe integration of an AM pressure reducer into a composite pressure vessel (CPV) was identified as an ideal platform to explore and evaluate the synergies between composites and AM for the following main reasons:Limitations of Conventional CPVs: major points of concern for users of pressure vessels are protruding pressure reducers (packaging, safety) and the disadvantages from missing integration of CPVs and reducers (weight, cost, etc.).Technical opportunity through combination of AM + composite technologies: AM offers advantages for complicated inner structures, such as pressure reducers, because complexity comes without extra cost and the process allows for designs that are otherwise not possible to manufacture.

CPVs are a structure ideally suited for composite materials due to their simple load case scenario and the resulting possibility of exploiting anisotropic features of fiber reinforcement.Market opportunity: CPVs are seeing a large market growth, based on two growth drivers:with the advancement of design and manufacturing technologies an increasing number ofpressure cylinders will be made from composite materials. the global demand for pressure vessels (composite and metal) is growing due to a sharp increase in usage, such as for automotive (hydrogen and CNG), leisure (paintball) and some niche markets (fuel-cells).5.2 Decision ProcedureThe decision procedure from Section 3 is applied to the composite pressure vessel application. To address the first two questions, these pressure vessels were already manufactured using filament winding. The issue is whether or not some level of integration can be achieved with external components.

The issue is whether or not some level of integration can be achieved with external components. When thinking about question 3, the idea of integrating the pressure reducer into an end-cap emerged as a possibility.To address question 4, it is important to realize that the end-caps were manufacturing using injection molding in polyamide 6. The liner was machined to size from a hollow tube of polyamide 6 polymer. Together they form the mandrel on which the body of the pressure vessel is fabricated, as shown in Fig. 2. Further, these components become parts of the final product, instead of utilizing hard tooling that is re- used to fabricate the next part. Taken together, questions 4 and 5 raise the possibility of consolidating pressure reducer functionality into an end-cap. The resulting design will contain geometric detail too small and complex for composites fabrication processes.

The resulting design will contain geometric detail too small and complex for composites fabrication processes. This complexity can be isolated in the end-caps, which will be fabricating using AM processes.When considering question 6, the cylindrical pressure vessel will continue to be filament wound. In fact, its design and manufacturing process remain largely unchanged from pressure vessel designs with simple, conventional end-caps. The end-caps with pressure reducer functionality will be fabricated by one of the polymer AM processes. Both powder bed fusion and material extrusion utilize polyamide materials, including polyamide 11 and 12, with glass-filled variants of these materials. These processes and materials will be leading candidates for the end-caps.2021Regarding question 7, polyamide 6 is used as the thermoplastic matrix for the pressure vessel, which is the same material as the end-caps and liner.

If polyamide 11 or 12 is used for the end-caps, the materials are very compatible. In fact, no additional joining considerations are needed for this example. This means that questions 8 and 9 can be addressed trivially.To address question 10, the existing filament winding process must be examined in more detail. As shown in Figure 2, spindle fixtures are bolted to the end-caps so that the spindle-mandrel assembly can be installed on the filament winding machine. This requires extensions on the end-caps that can accommodate bolts, which is readily achievable by adding material to lengthen the end-caps.

This requires extensions on the end-caps that can accommodate bolts, which is readily achievable by adding material to lengthen the end-caps. This extra material can be cut off after the pressure vessel is fabricated.As indicated by questions 6 and 11, it is important to consider functionality and economics of AMproduction of end-caps.After completing the decision process, the possibility of integrating pressure reducer functionalityinto end-caps and fabricating the end-cap designs via AM seems worthwhile to consider further.5.3 CPV DesignGiven the positive indications for proceeding from the decision process, attention can turn to redesigning the CPV with a focus on the end-caps. The design objective for the pressure reducer was to incorporate all of its components into the end-cap. It was deemed acceptable to increase the size of the end-cap, provided that its length was less than the total length of the original end-cap plus conventional pressure reducer.

It was deemed acceptable to increase the size of the end-cap, provided that its length was less than the total length of the original end-cap plus conventional pressure reducer. Also, pressure reduction capability was to be demonstrated, but only for “shop floor” pressure levels of up to 100 psi, rather than pressures to which the CPV is rated (~3000 psi). This enabled the exploration of design concepts and composites vs. AM trade-off exploration without having to perform time-consuming engineering studies.With regards to the design process proposed in Section 4, three geometric regions of the CPV were identified: the main cylindrical pressure vessel, one end-cap with the conventional design, and one end- cap with integrated pressure reducer. The design of the main CPV body remains unchanged, with typical filament winding patterns of PA6 thermoplastic matrix reinforced with continuous carbon fibers. Filament winding will be performed over a PA6 liner.

Filament winding will be performed over a PA6 liner. The plain end-cap will again be injection molded in PA6. That leaves the design of the end-cap with integrated pressure reducer, which will be described in the remainder of this subsection.After an initial conceptual design exploration in which over 10 design concepts were generated, several concepts were selected for more detailed design studies. One of these designs was developed to the prototyping stage, as shown in Fig. 4, and consists of one pressure reducing valve stage. A description of its operation will help with understanding the design. Gas from the pressure vessel enters the poppet chamber (region with the poppet spring), which tends to push the poppet forward into the seat, blocking the delivery of gas to the output.

Gas from the pressure vessel enters the poppet chamber (region with the poppet spring), which tends to push the poppet forward into the seat, blocking the delivery of gas to the output. If the user sets the valve handle to a positive pressure, the diaphragm spring compresses and exerts a force on the diaphragm which causes the poppet to displace and allows gas to pass into the diaphragm chamber. As pressure builds in the diaphragm chamber, the rising pressure exerts a force on the diaphragm that displaces the poppet back towards its original position, closing the valve. The balance of input and output pressures through the poppet displacement ensures that the output pressure stays at or below its setting.The design in Fig. 4 shows the pressure reducer stage adjacent to the outer boundary of the end-cap where the end-cap’s thickness is greatest.

4 shows the pressure reducer stage adjacent to the outer boundary of the end-cap where the end-cap’s thickness is greatest. Channels connect the pressure vessel to the poppet chamber through the feature labeled Restrictor and the diaphragm chamber to the output (labeled Output Air). In this design, the poppet and diaphragm were combined into a single part. This combined part and the part labeled Slider were to be fabricated separately, using AM, from the end-cap and assembled with poppet and diaphragm springs (purchased parts). Additionally, a cover part was designed to cover and seal the pressure reducer cavities.2022Valve["''"'· ReslllCIOIAl, comi11g f10111the 1est1lct0t behind the popet-s1>1 big assembtyPtessU1e-Redocl111.J V.llve Cil'lilyDlaph1a0tnCap of the Restrk:101l111Mn Handle Cil'lltyFig.

4: Initial end-cap design.To explore the capabilities of various fabrication processes, several versions were fabricated in the Stratasys J750 (a), Stratasys Fortus 450mc (b), EOS P396 laser sintering (c), and Markforged Mark 2 (d) printers, as shown in Fig. 5. End-caps from the J750 and P396 were of high enough quality to provide pressure reducing functionality. The material extrusion fabricated end-caps did not have a high enough feature resolution and precision to prevent leaks and had other assembly and operational problems. Since the materials used in the Stratasys J750 were acrylate-based photopolymers, it was decided to not attempt to use J750 fabricated end-caps for functional testing or fabrication of complete CPVs. However, the J750 was easy to use and resulted in transparent end-caps that enabled visualization of internal features, so was used to fabricate many end-cap prototypes.

However, the J750 was easy to use and resulted in transparent end-caps that enabled visualization of internal features, so was used to fabricate many end-cap prototypes. The material used in the P396 was polyamide 12 (PA12), which is the same family of thermoplastics used as the matrix material in the CPVs to be fabricated. Hence, PA12 end-caps fabricated in the P396 were selected for usage in the project demonstration CPV.a)b)c)d)Fig. 5: End-caps fabricated in a) Stratasys J750, b) Stratasys Fortus 450mc, c) EOS P396 laser sintering, and d) Markforged Mark 2 printers.Several months of engineering design, analysis, prototyping, and refinement were undertaken. Finite element analysis demonstrated that the end-cap thicknesses must be increased by approximately 20mm in order to prevent failure at 3000 psi pressure. With this additional space in the end-cap, the layout of the pressure reducer was changed with the objective of eliminating leaks.

With this additional space in the end-cap, the layout of the pressure reducer was changed with the objective of eliminating leaks. Two major design changes were2023made. First, the pressure reducer was placed inside the end-cap, without the need for a cover plate. This avoided significant sealing challenges. Second, an in-line configuration of chambers and components was used in order to facilitate assembly and sealing.The final end-cap design is shown in Fig. 6. The end-cap shown in Fig. 6b was fabricated in Vero Clear material on the J750 since its transparency allows visualization of internal features. For testing and filament winding, the end-cap was fabricated in polyamide 12 on the EOS P396. Internal components were fabricated using Vero White material on the J750 since they could be produced with good precision. The diaphragm and poppet spring holders were screwed into place.

Internal components were fabricated using Vero White material on the J750 since they could be produced with good precision. The diaphragm and poppet spring holders were screwed into place. O-rings and springs were sized appropriately for the pressure reducer functionality and were purchased as stock parts.C onnectinu line bet ween Two C avltiesOlffl>tffLeft Cy lindr ic'11 c,wityfo1· Olill)llrn!]lll Spring AssemblyC onnectin!] line From Pressure vessel t o Rigid Cavity~ulputFigure 6. Final end-cap design (a) and assembled prototype (b).After assembly, the final end-cap design was tested using a specially constructed pressure vessel andperformed successfully with air pressures up to 90 psi.20245.4 CPV ManufacturingTo fabricate the complete CPV, all of the components shown in Fig. 2a need to be fabricated and assembled. As mentioned, the liner is machined from stock pipe, and the standard end-cap is molded.

2a need to be fabricated and assembled. As mentioned, the liner is machined from stock pipe, and the standard end-cap is molded. Before either end-cap was fabricated, however, extensions were added to allow for mounting the end-cap and liner structure to a fixture for the process (Figure 7). The extension was designed as a sacrificial element to the end-cap, to be removed later in the manufacturing process. The fixtures on the ends of the assembly are standard components. These components were assembled in preparation for CPV filament winding.~ -1-LHi II I r•.LJ I L _____________ J +----+----+ :r: co11 II IFigure 7. The extension, indicated by EH (Extended Height), incorporated as mounting points for attachment fixtures, removable after completion of the winding process.Figure 8.

The extension, indicated by EH (Extended Height), incorporated as mounting points for attachment fixtures, removable after completion of the winding process.Figure 8. Laminate layup sequence and angles.The layup sequence for the composite vessel must be symmetric about its middle surface to reduce strain variations through the laminate wall and to avoid warping upon subjected to high internal pressures. The winding angle of the fibers is the major variable determining the mechanical performance of the pressure vessel. Based on references from literature [7], the optimum angle for a thin pipe vessel to withstand the maximum burst pressure is considered to be at ±55°. The laminate layup sequence and angles used are illustrated in Figure 8.Filament winding was performed with carbon fiber tapes impregnated with PA6 thermoplastic. Completed CPVs are shown in Figure 9. One pressure vessel with 3D printed end-caps was tested for leaks with air pressure up to 50 psi.

Completed CPVs are shown in Figure 9. One pressure vessel with 3D printed end-caps was tested for leaks with air pressure up to 50 psi. No leakage was evident in the joint regions between the pressure vessel and the end-caps.2025Figure 9. Fabricated CPV with AM-fabricated end-cap with integrated pressure reducer.6 ClosureThe issue of how to take advantage of the advantages of fiber-reinforced polymer composites and AM processes was explored in this paper. A decision guide and design process were proposed to guide the designer in designing a device that consists of both composite and AM parts. Overall, the guidance encourages the development of relatively large regions with simple shapes, and other, smaller regions with potentially complex shapes. The large, simple regions are candidates for composites, while the smaller, complex regions are candidates for AM.

The large, simple regions are candidates for composites, while the smaller, complex regions are candidates for AM. The decision guide and design process were illustrated by the design of a composite pressure vessel, where it was desired to integrate a pressure reducer into an end-cap of the vessel. The end-cap was redesigned to incorporate pressure reducer components and fabricated using powder bed fusion technology. Other pressure reducer components were designed and fabricated, including the poppet and diaphragm, and assembled with purchased springs and o-rings. After assembly, the AM fabricated end-cap performed successfully under air pressures up to 90 psi. The pressure vessel body was fabricated by filament winding with carbon fiber tapes impregnated with polyamide 6 thermoplastic.

The pressure vessel body was fabricated by filament winding with carbon fiber tapes impregnated with polyamide 6 thermoplastic. Several pressure vessels were fabricated with the redesigned AM end-caps and tested successfully for leaks.AcknowledgementThe authors acknowledge funding from the Singapore National Additive Manufacturing Innovation Cluster (NAMIC) from grant 2016260. Authors VH and DWR acknowledge usage of the Digital Manufacturing and Design Centre facilities at the Singapore University of Technology and Design.References[1] Weaver, P.M., “Designing Composite Structures: Lay-up Selection,” In: Proceedings of the Institutions of Mechanical Engineers Part G Journal of Aerospace Engineering, 216 (2002) 105-116.[2] Gibson I., Rosen D.W., Stucker B. (2015) Additive Manufacturing Technologies: 3D Printing, RapidPrototyping, and Direct Digital Manufacturing.

sign of a customized multi-directional layered deposition system based on part geometry.Prabhjot Singh, Yong-Mo Moon, Debasish Dutta, Sridhar KotaAbstractMulti-Direction Layered Deposition (MDLD) reduces the need for supports by depositing on a part along multiple directions. This requires the design of a new mechanism to re- orient the part, such that the deposition head can approach from different orientations. We present a customized compliant parallel kinematic machine design configured to deposit a set of part geometries. Relationships between the process planning for the MDLD of a part geometry and considerations in the design of the customized machine mechanism are illustrated. MDLD process planning is based on progressive part decomposition and kinematic machine design uses dual number algebra and screw theory.1. Introductionthe development of numerous Layered Manufacturing (LM) techniques.

Introductionthe development of numerous Layered Manufacturing (LM) techniques. The main advantage afforded by LM is its ability to produce geometrically complex parts without specialized tooling in a relatively short period of time. LM processes are characterized by the need for sacrificial structures to support overhanging regions of the part. This necessitates time consuming post- processing and degrades part quality.The past decade has seenMulti-Direction Layered Deposition (MDLD) Systems such as [3][4][5][7] build parts without supports by depositing a part along multiple directions. Depending on the process, either the deposition nozzle or the base table has multi-axis kinematics (refer Figure 1).(a)(b)Figure 1: (a) Deposition Table with rotational and translational actuation (b) Depositionnozzle mounted on a multi-axis robotic arm.266This necessitates a mechanism with more than 3 degrees of freedom.

A full 6 degree of freedom mechanism is expensive and is very often redundant for customers engaged in the design and fabrication of specific geometries. Recognizing this, recent research efforts such as [8][15] have proposed simplified kinematics for multi- direction deposition systems. While they have a distinct advantage over 2.5D LM systems, the fixed configuration of the nozzles/deposition table restricts support-less deposition to a few part geometries.In this research we propose a cost effective method to achieve multi-direction deposition by using kinematics which are customized to fabricate a specific family of parts.For a given geometric part family, our approach comprises of two tasks. The first task is the process planning underlying multi-direction layered deposition of a given member of the part family. The input part geometry is converted into sets of uniform parallel slices aligned along a set of vectors called the build directions.

The input part geometry is converted into sets of uniform parallel slices aligned along a set of vectors called the build directions. In the second task, slice geometry and build directions are used to derive the machine kinematics. Traditionally mechanism design has been a mixture of art and science. Researchers in the past have tried to classify, compile and codify basic elements of mechanisms in order to systematize the creative design process. In this research, we propose a compliant parallel kinematic machine (CPKM) design methodology [9][10]. The CPKM methodology synthesizes the mechanism using dual number algebra and screw theory. The synthesized mechanism has high accuracy because it uses compliant joints which are clearance and friction free mechanical component.

The synthesized mechanism has high accuracy because it uses compliant joints which are clearance and friction free mechanical component. A typical CPKM mechanism is shown in Figure 2.Figure 2: A deposition table constructed using the Compliant Parallel Machine DesignMethodologyThe remainder of the paper is organized as follows: Section 2 details the process planning methodology underlying the MDLD of a part. Section 3 presents a brief summary of the theory underlying CPKM and its relevance in the context of MDLD. In Section 4 we define an illustrative parametric part family along with its MDLD process planning. The output of the process planning information is used in267Section 5 to construct the kinematics of a CPKM. Section 6 concludes the paper with a summary of the research and directions for future research.2. Process Planning Methodology for MDLDA process planning framework for MDLD was proposed in an earlier publication [11].

Process Planning Methodology for MDLDA process planning framework for MDLD was proposed in an earlier publication [11]. It was assumed that the deposition nozzle is mounted on a multi- degree of freedom robotic arm. No other assumptions were made about the kinematics of the MDLD system. In [11] we addressed the core question in the analysis of multi- direction deposition systems, which is: how much of the part should be built in one direction and why? A brief overview of the process planning tasks is covered in this section. The reader is referred to [11] and [12] for further details.Assuming that the initial orientation of the part is user defined, the followingMDLD process planning tasks were identified:1. Decomposition of the part volume. 2. Establishing the Build Directions. 3.

Decomposition of the part volume. 2. Establishing the Build Directions. 3. Sequencing the Decomposed sub-volumes.The overall strategy for multi-direction slicing involves the progressive decomposition of the part (P) into sub-volumes each of which can be completely built along a certain direction. The input to each stage of this progressive decomposition is an unprocessed sub-volume (Vunproc) of the part which is processed using two operations:1. Find a build direction, B . 2. Using B to decompose Vunproc into buildable and unbuildable sub-volumes.The buildable part of Vunproc is classified as a processed sub-volume (Vproc) with B as its assigned build direction. The unbuildable sub-volume (Vunproc) forms the input to the next stage of the decomposition process. This stage-wise decomposition stops when:(cid:27)iVproc i(cid:32)( PVol)Where Vol(P) is the volume of the part P.2.1.

This stage-wise decomposition stops when:(cid:27)iVproc i(cid:32)( PVol)Where Vol(P) is the volume of the part P.2.1. Part Volume DecompositionThe overhang angle constraint as defined by Allen and Dutta [2] restricts the deposition of overhanging regions in LM. For a given build direction ( B ), the part volume decomposition algorithm is concerned with the identification and disjunction of part volumes which can be built and those which cannot. The algorithm defines surface regions to be unbuildable when the angle between the surface normals and the build direction is greater than 90 degrees. Such unbuildable surface regions are268bounded by silhouette edges of the surface [6] when viewed from – B , the opposite of the assigned build direction. Unbuildable part volumes are then identified by sweeping the unbuildable surface regions along B and subtracting the resulting swept volume from the part P. This is shown through an illustrative CAD model below (ref.

This is shown through an illustrative CAD model below (ref. Figure 3).BSilhouette EdgesP unbuildbuildP(a) Part Volume DecompositionPunbuild = VunprocBuild mapBuild map, discretizedSelected build direction = (0, 1, 0)(b) Build Direction DeterminationFigure 3: Part volume decomposition and build direction determination for MDLD.2.2. Build Direction DeterminationIn this task a build direction is assigned to an unprocessed sub-volume. It involves both the identification of all feasible build directions and the selection of the best build direction. Any vector which makes an angle of 90 degrees or lesser with all surface normals of the part boundary is considered to be a feasible build direction. The set of all feasible build directions is called the build map. The best build direction (a member of the build map) minimizes the average weighted cusp height as defined by Alexander and Dutta in [1].

The best build direction (a member of the build map) minimizes the average weighted cusp height as defined by Alexander and Dutta in [1]. An illustrative build map of a part volume and the selected build direction for a CAD model is shown in Figure 3b.2.3. Volume SequencingLayered deposition requires the presence of a base substrate. Using Figure 3a as a reference, Pbuild must be deposited before Punbuild since the base or the common face between the two sub-volumes lies on Pbuild. Consequently the deposition of Punbuild cannot269start till its base has been created, establishing Pbuild’s precedence in the deposition sequence.The overall process planning algorithm for MDLD is summed up in the following flowchart (ref. Figure 4). As mentioned, the orientation of the part relative to the deposition table, i.e. the initial build direction ( B ) is user defined.

Figure 4). As mentioned, the orientation of the part relative to the deposition table, i.e. the initial build direction ( B ) is user defined. The build direction determination module and volume decomposition are applied recursively till the entire part is decomposed. These are succeeded by Part volume sequencing and slicing along assigned build directions ( B ).Input : P, B = (0, 0, 1)(Pbuild, Punbuild)=Decompose(P, B )Sequence PbuildB = Build(Punbuild, B ) P = PunbuildSlice(Pbuild, B )Figure 4: Overall algorithm for MDLD process planningAn illustrative CAD model with its volume decomposition is shown in Figure 4a. The model is consequently sliced as shown in Figure 4b. The initial build direction is chosen along the Z+axis.(a)(b)Figure 3: (a) A CAD model with its MDS decomposition and associated build directions (b)Slices of MDS subvolumes along selected build directions.2703.

Mechanism SynthesisA machine for MDLD is comprised of two parts; 1) X-Y printing overhang head unit, and 2) the workpiece orientation unit. Generic LM machines have an X-Y printing unit that has coordinated control in the X-Y direction and a separate control for the Z direction. To design the MDLM machine, we need to add the workpiece orientation unit that aligns the building direction to the Z axis of the head unit. This paper presents the design of the workpiece support unit that is using the compliant mechanisms.Traditional mechanisms attain motion through the use of rigid links and discrete joints. They are ubiquitous in the world of machines and have been studied for centuries. Within the last several decades, there has been a growing interest in what have been termed compliant mechanisms. Unlike their rigid body counterparts, compliant mechanisms utilize the flexibility of their members to transmit or transform motion and forces.

Unlike their rigid body counterparts, compliant mechanisms utilize the flexibility of their members to transmit or transform motion and forces. While traditional mechanisms are designed to be stiff and strong, compliant mechanisms are designed to be flexible and strong.Not every problem is best solved with compliant mechanisms, but for applications better suited for compliant mechanisms, there are a number of significant benefits. They can be summarized as follows:(cid:131) Assembly: Compliant mechanisms are designed to be monolithic. Thus, there is a reduced need for mechanism assembly. In most cases, compliant mechanisms are designed to be coupled with actuators. This typically requires some assembly, but the number of parts is far less in comparison to traditional mechanisms.(cid:131) No wear: Compliant mechanisms attain mobility through deformation, and therefore do not experience wear. As such, there is a reduced need for maintenance such as lubrication.

As such, there is a reduced need for maintenance such as lubrication. Also, failure occurs either from static or fatigue failure. These types of failure are more predictable than wear phenomena.(cid:131) No backlash: Due to the absence of discrete joints, compliant mechanisms do notsuffer from backlash. As a result, high precision may be attained.(cid:131) Energy storage: Compliant mechanisms store elastic energy as they deform. This energy may be used to assist in applications requiring a return stage. There is a reduced need for springs and possible actuation.The design of multi-degrees of freedom mechanism is not always been a simple task while maintaining the accuracy and stiffness of the mechanisms. One of the most popular design of spatial multi degrees of freedom mechanisms is parallel kinematic machines (PKM) which is know to have high stiffness.

One of the most popular design of spatial multi degrees of freedom mechanisms is parallel kinematic machines (PKM) which is know to have high stiffness. However, PKM has poor accuracy compare to the serial mechanisms and planar mechanisms since it needs more mechanical joints than the other mechanisms. In this paper, we are using the compliant joint [9] based parallel kinematic machines; CPKM (Compliant Parallel Kinematic Machines) so the mechanism has high stiffness without the accuracy problems.The mechanism does not need the full six degrees of freedom since CPKM is customized around a parametric part family. In this paper, we take the building block approach to design a CPKM with less than six degrees of freedom [10].271CPKM is comprised of a top plate (end effector or work holding), a bottom plate, a constraining leg and active legs. The top plate holds the workpiece and the bottom plate is attached to the MDLM machine.

The top plate holds the workpiece and the bottom plate is attached to the MDLM machine. The constraining leg is a one serial kinematic chain that connects those two plates and defines the total degrees of freedom of the workpiece supporting unit. The active legs are attached after the constraining leg is configured so that the machine can be activated.To design CPKM, its requirements should be clarified first. The requirements for a CPKM are the required motions. To capture the characteristics of the motions, screw theory based dual vector representation and its algebra has been used in this research. When we think of a motion, it includes many aspects of the motion such as range of motion, type of motion, direction of motion and location of motion. The dual vector representation is one of the most suitable methods to represent these in explicit form.

The dual vector representation is one of the most suitable methods to represent these in explicit form. The algebraic operations and modeling methods used in this research is following reference [9].The format of the dual number representation of a motion (or screw) has four parts; 1) magnitude, 2) dual pitch, 3) direction vector and 4) coupled vector. The magnitude and dual pitch capture the required range and the types of the motions respectively. The direction vector and the coupled vector is a line represented in dual vector form, thus it contains the orientation and location of the motion axis.$ˆ (cid:32)ˆˆ LPM(cid:32) (cid:62) MM max,current,Mmin(cid:64)(cid:11)P A(cid:14)P (cid:72) L(cid:96)CD (cid:12)(cid:94) (cid:14) (cid:72)Where Mmax, Mcurrent, and Mmin represent the maximum, current and minimum value of the magnitude respectively. PA and PL represent the angular and linear pitches respectively.

PA and PL represent the angular and linear pitches respectively. And D and C represents the direction and couple vector of the lines.The range of motion is calculated as the Mmax – Mmin. If the magnitude is a single number then the motion is considered as the displacement from 0 to the magnitude. The dual pitch has two numbers; PA and PL. For a pure rotational motion and a pure translational motion, the dual pitches are (1 + (cid:72) 0) and (0 + (cid:72) 1) respectively. For a generic screw motion, the dual pitch is (1 + (cid:72) h) and the unit of the magnitude is angle. The direction vector (D) indicates the direction of the motion axis. The couple vector is the cross product of the location of the line and the direction vector (C = R X D). Therefore, the dual vector satisfies the following conditions.D(cid:32)1CD(cid:32)(cid:152)0For the multi-DoF motions, the dual vector itself contains variables that capture the DoF.

Therefore, the dual vector satisfies the following conditions.D(cid:32)1CD(cid:32)(cid:152)0For the multi-DoF motions, the dual vector itself contains variables that capture the DoF. However, it is not desirable to have more than one DoF for a motion. Therefore,272(1)(2)we will use multi-DoF dual vectors only for the intermediate motions and they will be decomposed into single DoF in the end.The decomposition of the motion is required to match the required motions to the kinematic structure of the constraining leg. The decomposed motions can be directly used to synthesize the constraining leg since the order of screws also determines the dependency between the motions as in a serial kinematic chain.The second step toward the design of CPKM is to configure the constraining legfor the required motions.Figure 6. CPKM Design processAfter configuring the constraining leg, the active legs should be added to complete the CPKM design.

CPKM Design processAfter configuring the constraining leg, the active legs should be added to complete the CPKM design. The generic design of the active legs is designed so that it has six DoF. However, the actual DoF of the active leg is five since the actuator constraints one DoF from the configuration. Now the active legs have five DoF, in other words each of them constraints one DoF, the active legs control the DoF of the CPKM with those constrained DoF and hence causes the singularity. In adding active legs, the arrangement should be carefully done otherwise it will cause redundancy or singularity.Figure 7. active leg design2734. Illustrative Part Family (PF) Definition and MDLD Process PlanningAccording to Shah and Mantyla [13] a solid model in which the main model entities, such as faces, edges etc., are related by geometric constraints defines a parametric solid model.

The geometric constraints specify mathematical relationships between the numerical variables of the model entities. For the purposes of this research we shall define a part family to be the set of parametric solid models which share the same topology and geometric constraints. A member of the part family is instanced by assigning a value to each of the geometric constraints. We do not concern ourselves with constraint satisfaction or constraint propagation as addressed in [13] and [14]. Instead, our objective is to show the relationship between MDLD process planning and CPKM synthesis in the context of geometrically varying solid models.An illustrative part family is shown in the following figure (ref.

MDLD Process Planning for PFMDLD process planning tasks for the part family are derived from those discussed in Section 2.0. The build directions for the part family vary with the geometric constraints which effects the resultant part volume decomposition. The process planning tasks for an illustrative part family member are shown in Figure 5. These include part volume decomposition and slicing. Uniform slice thickness is assumed throughout. Note that the sequence of deposition is V1 – V2 – V3, based on the existence of the base surface (as discussed in Section 2.0). The deposition nozzle is assumed to be oriented along the build direction.V3B2B3V2V1B1zy(a) MDLD decomposition(b) Multi-Directional SlicingFigure 5: MDLD Process Planning for an illustrative PF memberB1 ,(refer Figure 5b) the initial build direction is along the Z axis.

The following table summarizes the variation in the build directions, B2 and B3 with the variation in (cid:68) and (cid:69).Build Direction B1 B2Dependence (x, y, z) (cos(cid:68), 0, sin(cid:68)) (-cos((cid:69)(cid:3)–(cid:3)(cid:68)), 0, sin((cid:69)(cid:3)–(cid:68)))5. Compliant Mechanism for fabricating the PFThe example has two build directions, B1 and B2. The task of the workpiece orientation unit is to orient the workpiece from the up-right to the oriented angle of B2.275xTherefore one can assume that the mechanism requires one degree of freedom which is: rotation about the Z axis in the global coordinates.Table 1. Library of constraining legs for DoF requirementsFrom Table 1, the one rotation and no translation configuration is selected as the candidate configuration of the constraining leg.

Library of constraining legs for DoF requirementsFrom Table 1, the one rotation and no translation configuration is selected as the candidate configuration of the constraining leg. Since the mechanism need to orient the workpiece by angle (cid:68), the sizing of the compliant joint is determined to satisfy the motion range requirement [9].After having the constraining leg configured, the active leg is placed. Since the mechanism needs only one degree of freedom, it needs one active leg. To balance the active leg layout, we put two active legs as in the figure 6.276Figure 6: MDLD machine for PF memberAs the part family changes, the reconfigurable work support unit could be replaced for the new requirements.6. Conclusions and Future WorkIn this research we have presented a synergistic framework in which the kinematics of an LM machine are designed using the output from the MDLD process planning for a parametric part family.

The presented approach has the potential to be more cost effective when compared to MDLD systems which employ 5-6 axis kinematics. Our assumptions regarding the mode of material deposition are generic. Future research will be directed at interfacing with a specific deposition technique (such as Direct Metal Deposition (DMD) [7]). We will also focus on complex application specific part families. This will entail the inclusion of collision detection techniques resulting in the possible requirement for support structures and changes in the orientation of the deposition nozzle orientation.References[1] Alexander, P. and Dutta, D., “Part Orientation and Build Cost Determination for Layered Manufacturing”, Computer-Aided Design, 30, 1998, pp 343-358.277[2] Allen, S. and Dutta, D., “Wall Thickness Control in Layered Manufacturing for Surfaces with Closed Slices”, Computational Geometry: Theory and Applications, Vol.

5, pp. 55-60.[8] McMains, S., “Double Sided Layered Manufacturing”, Japan-USA Symposium on Flexible Automation, , Hiroshima, Japan , July 2002, pp.269-272.[9] Moon, Y. and Kota, S., “Design of Compliant Parallel Kinematic Machines”, AMSE DETC, Montreal, Canada, 2002.[10] Moon, Y. and Kota, S., “Automated Synthesis of Mechanisms using Dual-Vector Algebra”, Mechanism and Machine Theory, 37, 2002, pp 143-166.[11] Singh, P. and Dutta, D., “Multi-Direction Slicing for Layered Manufacturing”, Journal of Computing and Information Science in Engineering, June 2001.[12] Singh, P. and Dutta, D., "Offset Slices for Layered Manufacturing", ASME DETC, Montreal, Canada, September, 2002.[13] Shah, J. and Mantyla, M., Parametric and Feature-Based CAD/CAM, John Wiley and Sons, Inc., 1995.[14] Shapiro, V. and Vossler, D., “What is a Parametric Family of Solid?”, Third ACM/IEEE Symposium on Solid Modeling and Applications, Salt Lake City, 1995.[15] Yang, Y., Fuh, J., Loh, H.

and Vossler, D., “What is a Parametric Family of Solid?”, Third ACM/IEEE Symposium on Solid Modeling and Applications, Salt Lake City, 1995.[15] Yang, Y., Fuh, J., Loh, H. and Wong, Y., “Multi-orientation Deposition for Support- less Layered Manufacturing Process”, preprint.278.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesign of a Desktop Wire-feed Prototyping MachineYu-Chuen Chang*, Richard H Crawford*Department of Mechanical Engineering, The University of Texas at Austin, Austin, TX 78712AbstractMuch additive manufacturing research focuses on systems suitable for industrialapplications, especially research on metal processing. Our research aims to design a desktop-scaleprototyping machine to process metal wire. Possible applications for this research include variouswire structures, such as wire sculptures.The wire joining technology is the most important subsystem of the envisioned layer-basedprocess. In this research, three concepts are proposed and analyzed according to the power requiredto fully melt the wire.

In this research, three concepts are proposed and analyzed according to the power requiredto fully melt the wire. The selected approach uses a wire bender to create the desired geometry ofthe product, and a pulse TIG welder to join the metal wire to retain the shape. Experiments wereconducted to evaluate the joining strength of pulse TIG welds to verify the joining efficiency ofthe method. The experimental results indicate that filler metal is required to produce acceptablewelding strength. A conceptual CAD model of the complete system is presented.IntroductionThe appearance of desktop 3D printers has boosted the trend toward personal fabrication or“making” [1]. These desktop layer-based machines provide the ability for users to automatically create components with complex geometry. They not only benefit makers but also engineeringeducation.

These desktop layer-based machines provide the ability for users to automatically create components with complex geometry. They not only benefit makers but also engineeringeducation. However, most desktop machines are suitable for thermoplastics only, not for metals.One limitation for the development of a desktop metal prototyping machine is due to availableenergy. It is common for additive manufacturing technology to use a heat source, such as a laser,to soften or join precursor material. A comparatively large amount of power is required to completea metal joining process relative to polymer joining.Recent research and new products [2-7] are focused on processing metal material with layer-based approaches. These fall into three primary categories. The first category usesmodified heating1705processes. Unlike traditional metal additive manufacturing processes that use a laser, such asSelective Laser Sintering (SLS), Anzalone et. al.

The first category usesmodified heating1705processes. Unlike traditional metal additive manufacturing processes that use a laser, such asSelective Laser Sintering (SLS), Anzalone et. al. develop a 3D printer with gas-metal arc welding(GMAW) to fuse and deposit carbon steel wire (ER 70S-6) [2]. Another example is the 888 printer(Ability 3D, Casselberry, FL, USA), which integrates a welder and a router. The 888 printer canprocess aluminum and steel [3]. The second category uses non-traditional precursor material, suchas ColorFabb CopperFill metal filament [4] and Liquid Metal Jetting (LMJ), which uses liquidmetal as the pre-cursor material [5]. The third category separates the geometry-forming processand the heating/joining process. The Mini Metal Maker uses metal clay, a material containingmetal particles and binder, to create green components. The green component is then heated in akiln to 600-900C to remove the binder [6].

The green component is then heated in akiln to 600-900C to remove the binder [6]. Selective Inhibition Sintering (SIS) extrudes an inhibitor to create the contour of the geometry. The material powder blank is sintered in a furnace,and the inhibitor is the removed by water or an organic solution [7].The purpose of our research is to develop a different approach to manufacturing with metalwire. The research focuses on applications such as wire sculptures (See Figure 1). A commonmaterial for wire sculptures is copper, which is considered as one target material for our process.Figure 1. Example of wire sculptors [8]Functional Model of Additive Manufacturing1706Additive manufacturing processes include five primary subsystems: signal control system,material delivery system, geometry forming process, material joining process, and support. Figure2 shows a common functional model of additive manufacturing processes.Williams et. al.

Figure2 shows a common functional model of additive manufacturing processes.Williams et. al. [9] developed a functional framework for the conceptual design of additivemanufacturing technologies based on physical principles. The sub-functions identified by theminclude store material, pattern material, pattern energy, create primitive, provide new material, andsupport previously deposited material. In our research, we assume metal wire is processed duringthe additive manufacturing process. Therefore, the framework from Figure 2 is recast as a functionstructure in Figure 3.Figure 2. Framework of the additive manufacturing processesFigure 3. Function structure of additive manufacturing processesThe sub-functions used in our research are transport material, position material, shape material, couple material, and stabilize material. Note that “separate material” demonstrates the process of1707separating the pre-cursor material from printed material.

Note that “separate material” demonstrates the process of1707separating the pre-cursor material from printed material. Table 1 compares sub-functions identifiedin our research and and those proposed in [9].From [9]This ResearchStore MaterialWirePattern MaterialTransport MaterialPattern EnergyCouple MaterialCreate PrimitiveShape MaterialSeparate MaterialProvide New MaterialTransport MaterialSupportStabilize MaterialTable 1. Comparison of sub-functions between this research and [9]Conceptual DesignsA morphological matrix [10] was developed based on the six sub-functions in Figure 3.

Comparison of sub-functions between this research and [9]Conceptual DesignsA morphological matrix [10] was developed based on the six sub-functions in Figure 3. Table2 illustrates the morphological matrix with three conceptual designs: bending, laser, and welding.Sub-functionBending ConceptLaser ConceptWelding ConceptTransport MaterialRollerRollerRollerPosition MaterialMoving PlatformMoving ExtrusionMoving PlatformShape MaterialWire BenderLaserInduction WeldingCouple MaterialPulse TIG WeldingLaserInduction WeldingSeparate MaterialPulse TIG Weldingn/an/aStabilize MaterialTrussTrussTrussTable 2 Concept variations based on the morphological matrix Bending ConfigurationIn this concept, a wire bender is used to create the geometry of the component. Then, theshaped wire is joined by a welder.

Then, theshaped wire is joined by a welder. Since the geometry of the wire is created by the bender, itis possible to limit joining to specific spots (join discontinuously) instead of join continuously. Laser ConfigurationThe laser configuration uses a desktop laser system to fully melt the metal wire continuously. Possible choices for the laser system include CO2 laser and Nd:YAG lasers. Cost is an important consideration when selecting a commercial desktop laser system.1708 Welding ConfigurationThe welding concept uses a traditional welding process to melt the metal wire. Possiblechoices for the welding process include Gas Metal Arc Welding (GMAW), Tungsten InertGas (TIG) Welding, and induction welding. Since GMAW and TIG welding have beenstudied [2,3], this research focuses more on evaluating induction welding.Concept EvaluationConcept evaluation was conducted in order to investigate the feasibility and comparativeperformance of each concept.

The evaluations of the laser and the induction weldingconfigurations were based on simulations, and the bender concept was investigated by experiment.Continuous Joining and Discontinuous JoiningIt is important to evaluate whether the precursor material can be fully joined with a reasonablepower input. Most additive manufacturing processes are continuous joining processes; that is, thepattern energy is delivered continuously to combine precursor material. Here, discontinuousjoining is explored for the bending concept. Since the geometry is formed by a wire bender, it isonly necessary to join selected points to create a stable part. Therefore, a pulsed joining methodwith comparatively low power requirement can be used in the discontinuous joining process.

Therefore, a pulsed joining methodwith comparatively low power requirement can be used in the discontinuous joining process. Thissignificantly reduces the required power for the process, as shown below.Laser Concept EvaluationTo evaluate the joining efficiency of the laser, it is important to investigate whether the metalwire can fully melted at selected reasonable speed. As stated earlier, this evaluation was conductedby simulation. The material properties for the simulation are provided in Table 4.

The material properties for the simulation are provided in Table 4. Equation (4.1) is a heat transfer model developed to validate the laser continuous joining process.MaterialAluminumCopperIron𝒌 (𝐰 𝐦𝐊23740180.2) 𝑪𝒑 (𝐉 𝐤𝐠𝐊903385447𝐤𝐠 𝐦𝟑) 2702) 𝝆 (89337870𝐦𝟐 𝐬 97.1 × 10−6 117 × 10−6 23.1 × 10−6𝜶 ()𝑻𝒎(𝐊)93313581810Table 4 Material properties for the simulation1709∇2𝑇 +1𝑘𝑞𝛿(𝑥 − 𝑣𝑡)𝛿(𝑦 − 0)𝛿(𝑧 − 0) =1𝛼𝜕𝑇𝜕𝑡where 𝑇 = transient temperature (K),𝑘 = conduction coefficient (WmK⁄),𝑞 = rate of energy production by a specific heat source (W), 𝑣 = constant velocity of heat source (ms⁄ ),𝛼 = diffusion coefficient (ms2⁄ ), 𝛿(𝑥) = Dirac delta function, and 𝑥, 𝑦, 𝑧 = Distance from the heat source in x, y, and z directions.The three-dimensional model can be converted into a one-dimensional model since theprimary concern is heat conduction within the copper wire.

Additionally, the model is assumed tobe quasi-steady state. By substituting ε = x  vt. Equation (4.1) can be rewritten as∂2𝑇 ∂ε2 +1𝑘𝑞,,𝛿(ε) = −𝑣𝛼𝜕𝑇𝜕εThe final solution is𝑇 = 𝑇0 +𝑞,, 𝑣𝜌𝐶𝑝, 𝑥 < 0𝑇 = 𝑇0 +𝑞,, 𝑣𝜌𝐶𝑝𝑒−𝑣 𝛼𝑥, 𝑥 ≥ 0The goal of the simulation is to fully melt 0.64 mm metal wire, which is a common wiregauge for handcrafts [17]. The input heat flux of the laser can be defined by equation (4.4). Thespot diameter is assumed to be 0.4 mm in the simulation [16]. Table 5 shows the reflectivity of materials under CO2 laser and Nd:YAG laser irradiation. The relationship between speed (10 mm/s to 20 mm/s) and laser power is demonstrated in Figure 3 (CO2 laser) and Figure 4 (Nd:YAG laser).𝐼 = (1 − 𝑅)𝑃𝐴where 𝐼 = power intensity (Wm2⁄),𝑅 = reflectivity,1710(4.1)(4.2)(4.3)(4.4)𝑃 = laser power (W), and A = area of spot size (m2).CO2 laserNd:YAG laserAluminum0.940.8Copper0.970.96Iron0.880.68Table 5.

Reflectivity of various materials under CO2 laser and Nd:YAG laser [11]Figure 3. Power requirement for aluminum, copper, and iron for various scanning speeds under CO2 laserFigure 4. Power requirement for aluminum, copper, and iron for various scanning speeds under Nd:YAG laser1711Figure 3 and Figure 4 show the power requirement to melt copper is much higher than foraluminum and iron. The laser concept is potentially feasible for both iron and aluminum, especiallyfor a Nd:YAG laser.Induction Welding Concept EvaluationThe minimum required power for the induction welding concept can be evaluated by equation(4.3). Figure 5 shows the relationship between power and velocity (10 mm/s to 25 mm/s). Theprice of a 5000 W induction welder, such as the UltraFlex UPT S5/2000 (Ronkonkoma, NY, USA),is approximately 10,000 USD, which is in a similar price range as a 45 W laser system with acooling system [12].Figure 5.

Minimum induction welding power for various velocitiesBending Concept EvaluationThe bending concept includes a discontinuous joining process. A TIG pulse welder wasevaluated as a feasible joining technology. For this research, we evaluated the DingXing DX- YQ0505 TIG pulse welder with 120 V power input using a 2 % lanthanated electrode with 1/16”diameter and 38 tip angle. The peak output current is approximately 2828A over 5 ms, and the output arc voltage is 161V.An Instron Model 2519-107 tensile test machine was used to determine the breaking force ofthe weld in order to evaluate the joining performance. The specimen material was annealed copper(ASTM 3B, 22 AWG). The gauge length was 30 mm. The specimens were straightened andcleaned with acetone before butt welding. The speed of the tensile test was 1mm/min at room temperature. The details of the tensile test are outlined in ASTM E8 [13].1712Figure 5 shows the results of the tensile tests of the spot welds.

The speed of the tensile test was 1mm/min at room temperature. The details of the tensile test are outlined in ASTM E8 [13].1712Figure 5 shows the results of the tensile tests of the spot welds. The stress-strain curve ofannealed copper wire is given as a reference (See Figure 6). The breaking point of the weldingspecimens (20-40 N) is much less than the breaking force (56N) of the annealed copper wire.Additionally, the strain of the welded specimens is smaller. The primary reason is due to themisalignment of the welder and anisotropy of the welded joint. These two factors cause thebreaking force for the pulse welded joint to be far less than that of the copper wire. Figure 6 alsocompares the copper wire with a filler metal, EURO TOOL SS75 (a typical chemical compositionhard silver solder is 75% Ag, 22% Cu, and 3% Zn). The result shows that the wire breaks in otherlocations and not at the joints.

The result shows that the wire breaks in otherlocations and not at the joints. This demonstrates that the use of a filler metal can improve the jointstrength.Figure 5 Tensile test for welding copper wire1713Figure 6. Tensile test result for the copper wire and brazing with filler metalDiscussionThe laser and bending concepts show potential to produce additively manufacturedcomponents. A low reflectivity material is suitable for the laser concept. However, the main issuefor the laser is cost. A cooling system is required for the laser concept, and the cost for the laserconcept is at least $6000 [12]. On the other hand, the bending concept shows potential to deal withhigh reflectivity metal. The primary issue in this research is that the breaking force of the weldingspecimen is less than 56 N, which is easy to damage.

The primary issue in this research is that the breaking force of the weldingspecimen is less than 56 N, which is easy to damage. Filler material is required to improve thestrength of the welding joint.The bending concept is selected for further development because high reflectivity material iscommon for jewelry or wire sculptures (copper).Design of the Wire-Feed Prototyping MachineThe bending concept includes a pulse TIG welding system and a wire bender. A compressionbender is selected because it is suitable for the desktop design. The detailed components and sub-functions of the bender are shown in Figure 7 and Table 6.Figure 7. Function structure for a wire bender [14]The proposed CAD model for the desktop metal wire-feed prototyping machine is shown inFigure 8. The wire bender is designed to operated\ in the vertical direction to easily position theshaped wire. The welder torch can rotate to cut the wire or weld the wire.

The wire bender is designed to operated\ in the vertical direction to easily position theshaped wire. The welder torch can rotate to cut the wire or weld the wire. For annealed copper wire, the recommended average energy set up to cut 0.64 mm diameter wireis at least 8000 J basedon experiments [14].1714Sub-functionComponentImport WireMotor-driven wire feederTransport WireRoller setsStraighten WireRoller setsGuide WirePlastic pipeBender WireBend die, sliding pieceSeparate WirePulse TIG welderExport WireMoving platformTable 6. Sub-functions and components of the wire bender for the metal prototyping machineFigure 8. Proposed structure for the metal wire-feed prototyping machine [14]Conclusion and Future WorkThis research proposes a novel design for a desktop metal wire-feed additive manufacturingmachine.

Proposed structure for the metal wire-feed prototyping machine [14]Conclusion and Future WorkThis research proposes a novel design for a desktop metal wire-feed additive manufacturingmachine. Topics for further development include:1715Improve the welding strengthBased on Figure 5, the welding strength is not high and is non-repeatable. It is necessary toinvestigate the strength of the joint when filler material is used in the process. Develop the three-dimensional joining processIn this research, the design focuses on two-dimensional structures. Three-dimensionalstructures such as lattice structures [15] require further studies. Additionally, another controlchannel may be required in order to accurately position the shaped wire for three-dimensionalstructures. Integration of the system is necessary to investigate these questions. Modify the control programSoftware development is required.

SIGN OF A GRADED CELLULAR STRUCTURE FOR AN ACETABULAR HIP REPLACEMENT COMPONENTHongqing Vincent Wang a R&D EngineerScott R. Johnston b Postdoctoral FellowDavid W. Rosen b* Professora IronCAD Inc., 700 Galleria Parkway, Suite 300, Atlanta, GA 30339 b The George W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology Atlanta, GA 30332Corresponding Author: 404-894-9668, david.rosen@me.gatech.edu.ABSTRACTReviewed, accepted September 14, 2006The state-of-the-art porous coatings become more and more popular in uncemented prostheses to make bone grow into implants for biological fixation. In this paper, graded cellular structures are proposed for uncemented prostheses to enhance stability on implant-bone interfaces. As an example study, the authors develop a new acetabular implant with gradient porosity for hip replacement. A gradient porous acetabular component with cellular structure could match the bone’s elasticity.

A gradient porous acetabular component with cellular structure could match the bone’s elasticity. Material is adaptively distributed from high porosity at the bone-implant interface to solid metal at the joint’s articulating surface. The new acetabular prosthesis would replace metal-on-polyethylene bearing with metal-on-metal bearing for less wear. The design problem of acetabular component is formulated and a requirement list is elaborated. A detailed design of the prosthesis with a graded cellular structure is presented. The design concept is validated with a comparison to the existing products according to the design requirements.KEY WORDSCellular Structure, Unit Truss, Gradient Porosity, Implant-bone Interface, Hip Replacement,Uncemented Prosthesis, Acetabular Implant1INTRODUCTIONJoint replacements with prosthesis are becoming more common as the population of the world begins to age.

One major issue of prosthesis design is the fixation between the implants (commonly metals) and their surrounding bones. Uncemented prostheses emerged as a new technology to biologically fix implants to bones without the use of cement. Recent clinical results show that uncemented prostheses reduce the aseptic loosening rate and thigh pain in hip replacement [1]. The uncemented prostheses use porous metal surfaces to allow bone ingrowth into the surrounding prostheses as shown in Figure 1 [2, 3]. The uncemented prostheses are coated with multiple layers of micro-porous metal [1, 2] The bone ingrowth into the coating provides additional bonding strength to hold the implant in the desired position.Researchers have developed various porous coating techniques for biological ingrowth. Typical porous surfaces are plasma sprayed surfaces, sintered beaded surfaces with large spheres, sintered beaded surfaces with small spheres, and diffusion-bonded fiber-metal surfaces [4].

Typical porous surfaces are plasma sprayed surfaces, sintered beaded surfaces with large spheres, sintered beaded surfaces with small spheres, and diffusion-bonded fiber-metal surfaces [4]. Popular metal coating materials are titanium and tantalum. Common porous coatings have a volume porosity111between 30-70%, with an average pore size ranging between 100 and 1000 m(cid:80) , and a coating thickness between 500 and 1500 m(cid:80) [5]. The experimental data shows that a mean pore size 200 m(cid:80)(cid:116)and porosity 40%(cid:116)are experimentally optimal for bone ingrowth [4].The current coating techniques provide arguably sufficient bone ingrowth to enhance the implant fixation. The fatigue failure on the implant-bone interface has not been significantly reduced due to a sharp difference among the elastic modulii of the solid metal implant, coating, and bone.

The fatigue failure on the implant-bone interface has not been significantly reduced due to a sharp difference among the elastic modulii of the solid metal implant, coating, and bone. The coating porosity and thickness may not be the best for bone ingrowth since most of the implant is solid and has a higher effective density. The porous coating is uniformly applied onto the implant and the thickness cannot exceed 1500 m(cid:80) . The interconnectivity in the coating is limited because the only currently available coating particle geometries are spheres, beads, and diffused fibers [4].A graded cellular structure can provide an adaptive and ordered porosity distribution. As displayed in Figure 2, the graded cellular structure gradually can change from thin struts with high porosity and small elasticity on the exterior boundary to thick struts with low porosity and large elasticity on the inner boundary.

In this paper, the authors propose to use graded cellular structures in prostheses to enhance joint stability after implantation.Figure 1. Uncemented metal-bone interface with porous metal coating in prosthesis [1, 2]Figure 2. An example of cylindrical graded cellular structures (top view)The bone ingrowth into the porous implant is greatly influenced by the chemical, mechanical and geometric characteristics of the porous coatings in the uncemented prostheses. An elastic modulus matches to the surrounding bone, larger pore spaces, higher porosity and interconnectivity, and improved surface bioactivity can significantly increase bone growth into the gap region and porous coating [6]. The implant with a matching elastic modulus to the bone has less damage on neighboring tissues and a higher fatigue resistance at the bone-implant interface [7].

The implant with a matching elastic modulus to the bone has less damage on neighboring tissues and a higher fatigue resistance at the bone-implant interface [7]. Higher porosity, higher interconnectivity, and larger pore spaces can improve the surface bioactivity with more breathability to enhance bone ingrowth for stronger prosthesis fixation [4, 8].The authors present an example of an acetabular component using graded cellular structures for hip replacement. The new prosthesis design can gradually change its elastic modulus, achieve high porosity and pore interconnectivity, and reduce wear rate at the implant-bone interface.1122 DESIGN REQUIREMENTS FOR ACETABULAR COMPONENT IN HIPREPLACEMENTA conventional acetabular component usually consists of a polyethylene liner and metal shell. In this research, the acetabular component is designed as a single hemispherical part shown in Figure 3 (measurements taken from CAD models supplied by [9]).

In this research, the acetabular component is designed as a single hemispherical part shown in Figure 3 (measurements taken from CAD models supplied by [9]). The measured dimensions of the femoral component, femur and Ilium, the inner and outer radii of the acetabular component are 13.89 mm and 29.50 mm respectively, and the resulting acetabular thickness is 15.61 mm. The gap between the implant and bone must be less than 0.5 mm to allow bone growth into the implant. The total geometric tolerance (resulting from design, manufacturing, and installation) of the hemispheric implant should be no more than 0.5mm, thus the outer radius should be 29.50±0.25 mm. The inner surface as the prosthesis’ articulating surface should be very smooth and is usually polished. The interior region of Ilium surrounding the acetabular component is spongy bone with an average elasticity of 1 GPa, and yield strength of 5 MPa.

Biocompatible with bone 3.2. Less or no thrombophlebitis or infection4. ManufacturabilityD W W4.1. Manufacturable 4.2. Less delivery time 4.3. Lower cost5. InstallationW5.1. Less or no assembly3 MATERIAL SELECTION AND MANUFACTURING PROCESSThe first design issue is the material selection for best biocompatibility and sufficient mechanical strength. Biocompatible polymer and ceramic scaffolds (such as polyethylene) have been extensively studied for prosthesis components [12]; however, the mechanical properties of the polymer scaffolds are insufficient to support bone growth under loading conditions, and the ceramic scaffolds are too brittle. A few metal scaffolds, such as titanium and tantalum foam can provide sufficient strength and fatigue resistance because they maintain sufficient mechanical strength and fatigue resistance over extended periods in vivo [13].

Porous titanium as well as porous titanium coating have been successful at encouraging bone ingrowth both in vivo and in clinical trials [4, 14]. Moreover, titanium possesses excellent biocompatibility with body tissues and is widely used as an implant material. Titanium has excellent fatigue and corrosion resistances. A typical titanium alloy, Ti-6Al- 4V, has a yield Strength 900 MPa, ultimate tensile strength 960 MPa, elasticity modulus 110 GPA, and hardness 35 Rc. However, it has a known drawback of its poor resistance to wear, and notch sensitivity when used as bearing material [16]. Tantalum carbide coating is a good candidate for protective hard coatings against wears in sliding applications; therefore, a layer of tantalum carbide coating (780±50nm thick and 53±4% carbon content) may be applied to the articulating surfaces to114reduce the wear [17].

The sliding wear resistance test on a conventional ball-on-disc test apparatus demonstrated that the friction coefficient of tantalum carbide (low carbon) coated surface can be as low as 0.15, which is comparable to the friction coefficient of untreated graphite, 0.14 [18]. In our design, titanium alloy has been selected as the material for cellular structure and solid liner of the acetabular component, and a layer of tantalum carbide will be applied as a coating on the liner’s articulating surface.With the development of additive manufacturing (AM) processes (also known as rapid prototyping), the manufacture of graded cellular structures becomes feasible. AM is capable to manufacture the entire acetabular component as a single part without assembly [19, 20].

AM is capable to manufacture the entire acetabular component as a single part without assembly [19, 20]. AM processes have the potential to manufacture the acetabular component including the porous structure and solid liner section as a single part without assembly as well as with varying architecture, porosity, and pore size of the cellular structure. Selective Laser Melting (SLM) is one of the possible processes to manufacture 3-D porous metallic structures using a variety of material options, including stainless steel, titanium, and chromium-cobalt [21, 22]. SLM resolution is capable of fabricating a strut diameter as small as 0.1mm.

SLM resolution is capable of fabricating a strut diameter as small as 0.1mm. The inner surface may be polished to an average roughness of 60nm for better smoothness and coated with a layer of tantalum carbide via vapor deposition [17].4 MATCHING ELASTIC MODULII WITH GRADED CELLULAR STRUCTUREA high porosity is required for adequate bone ingrowth into the implant to create adequate fixation to withstand normal physiological loading. Additionally, the implant’s mechanical properties should closely match the host’s tissue properties at the time of implantation [23, 24]. Since the acetabular component is surrounded by the acetabulum’s spongy bone (after implantation), the desired elasticity of exterior layer of cellular structure in the acetabular component should be 1 GPa.The microstructure of a graded cellular structure can be modeled as a unit truss with a desired elastic modulus value depending upon the truss diameter.

An octet unit truss structure has been chosen for this study and two octet trusses with different strut diameters are shown in Figure 4. The elastic at modulus of the implanted acetabular component is desired to change gradually from of the solid titanium. The effective elasticity of the unit GPa its boundary, (the Ilium bone) to hL denotes the half-strut length and d denotes the strut effE , is defined in Equation 1 [25] where truss, diameter. To match the desired elasticity at the exterior surface (the Ilium bone), the effective GPa as possible. Therefore, using Equation 2, the elasticityE 01.0(cid:32)sE110(cid:32)GPaeffE needs to be as close toE 0(cid:32)1.0calculated ratio between the strut diameter and half-strut length isd L h(cid:32)0.1241.

Porosity decreases as diameter increases5POROSITY AND PORE SIZE OF GRADED CELLULAR STRUCTUREA high degree of interconnected porosity can significantly help bone growth into the implant [13]. A porosity of 90% is recommended for optimum diffusive transport within a cell–scaffold construct under in vitro conditions [23]. The advantages of scaffold constructed with large surface area to volume ratio has been discussed and demonstrated [24].The material volume and surface area of the unit trusses with various strut diameters are measured from their solid models. The porosity and surface/volume ratio decreases when the ratiobetween the strut diameter and strut half-length increase, as displayed in Figure 6. Whend L h(cid:32)0.1241,the porosity can reach as high as 91.17% and create a section of the cellular structure having anexcellent diffusive transporting capability; conversely, the porosity approaches 0% whend L hisapproximately 0.65.

Relative surface/material volume ratio decreases as diameter increasesFigure 8. Pore size of octet trussBrandon has researched various metallic coatings used in hip implants and has determined that the total amount of bone ingrowth does not significantly differ for pore sizes ranging between 200 and 450 μm, but was significantly less for a pore size of 140 μm [4]. Therefore, the optimal pore size for hip implants should be greater than 200 μm. Unfortunately, there is a lacking in literature concerning the maximum allowable pore size.In our designed graded cellular structure of the acetabular component, the strut diameter starts from 100 μm at the outer surface of the cellular section.

A cross-section view of the STL model of the new acetabular component with graded cellular structure119Figure 11. The entire new acetabular component with graded cellular sectionFigure 12. Mechanics analysis using 1/12 of the new acetabular component6 MECHANICS ANALYSIS OF THE CELLULAR STRUCTURED ACETABULARCOMPONENT AND VALIDATION OF THE CONCEPTUAL DESIGNThe hip joint contact area and pressure distribution during activities of daily living vary from one movement and to another and are relatively complicated. Yoshida has determined that the peak pressure of moderate magnitude is located at the lateral roof of the acetabulum during fast, normal, and slow walking. The peak pressure is located at the edge of the posterior horn during standing up, sitting down, and knee bending motions [27].

The peak pressure is located at the edge of the posterior horn during standing up, sitting down, and knee bending motions [27]. The peak pressure varies from 2.87MPa during slow walking to 9.36MPa during sitting down.For simplicity, the authors have analyzed 1/12 of the acetabular component (shown in Figure 12) and applied a uniform pressure, 9.36MPa, onto the acetabular cup (titanium liner). The external surface is fixed in all 6 degrees of freedom. There are a total of 1255 nodes attached to the inner 149.4mm ) and each node bears an equivalent 1.1N force along the radial surface (with an area of direction of the hemispherical cup. The maximum stress among all struts is 23.6MPa with a maximum deflection among all nodes is 0.028mm . The total strain energy stored in the cellular structure is.

The maximum stress among all struts is 23.6MPa with a maximum deflection among all nodes is 0.028mm . The total strain energy stored in the cellular structure is . The effective elasticity of the cellular section in the radial direction is 3.81 GPa, while 0.163N mm(cid:152) the effective elasticity is 1.27 GPa at the outer surface and 23.9 GPa at the inner surface.2The design of new acetabular component is validated through a critical evaluation with the design requirements elaborated in Table-1 and a comparison with current designs is shown in Table-2. A “+” and “-” denote whether the new design is superior to the existing designs, while a “0” denotes no improvement over the current design. All the demand requirements (D) were satisfied and most of wish requirements (W) were accommodated with improved results.7 CONCLUSIONSThe authors successfully use a graded cellular structure in designing an uncemented acetabular prosthesis for enhanced stability on implant-bone interface.

The new design of the acetabular component matches the elasticity of the surrounding bone and gradually changes from high porosity at the implant-bone interface to solid metal at the liner section. Other issues, such as material selection,120manufacturing process, pore sizes, surface/volume ratio have been discussed. The interconnectivity of cellular structure is superior to the porous metal, which are commonly spherical particles or diffused fibers. The elasticity of cellular structure is reconfigurable to match bone’s elasticity, while the porous metal is uniform and determined by the manufacturing process and material. The new acetabular component has better fatigue strength and significantly less wear since the metal-on-metal bearing replaces the metal-on-polyethylene bearing.

The new acetabular component has better fatigue strength and significantly less wear since the metal-on-metal bearing replaces the metal-on-polyethylene bearing. Less assembly is required due to fewer parts.In order to get closer to transfer this technology to clinical application, several unanswered questions related to implant design must be addressed in the future before a transfer of this technology to clinical trials can be conducted. The architecture type (one of the morphology issue other than porosity and pore size), physical prototyping, part inspection method, and physical experiments for mechanical property analysis have yet to be currently resolved.Table-2 Critical evaluations on the designed acetabular componentD / WCriteria+ /-Comments: comparison with existing acetabular cup with porous coating1. GeometryD1.1.

GeometryD1.1. Shape fitness (inner radius: 13.89mm;outer radius: 29.50mm; gap between implant and bone 0.5mm (cid:31))Gap could be greatly influenced by the manufacturing (e.g. SLM), not as good as existing designs. But it could reach the requirement.W1.2. Cellular architecture+ Type of architecture was not discussed in this research, however capable to provide various architectures. Existing designs have limitation from coating techniques.WD1.3. Appropriate porosity ( 40%(cid:116)onimplant-bone interface; solid on joint articulating interface) 1.4. Pore size ( 0.2mm (cid:116); avoid space collapse and best interconnectivity)+ 90.17% at the implant-bone interface, solid in liner section. Superior to the existing porous coating.+ Satisfied: average pore size 1.1mm at the implant-boneinterface; good interconnectivity.2. Mechanical propertiesD2.1.

Biocompatible with bone0 Satisfied: perfect biocompatibility between titaniumW3.2. Less or no vein inflammation, infectionand tissue. Not evaluated yet.D4. Manufacturability 4.1. ManufacturableSatisfied: minimum manufacturable feature size can be 0.1mm by SLMW WW5.4.2. Less delivery time 4.3. Lower cost Installation 5.1. Less or no assembly0 A few days. - High cost of using rapid prototyping. But acceptable.+ The regular polyethylene liner is removed. Lessassembly required.121ACKNOWLEDGEMENTSWe gratefully acknowledge the U.S. National Science Foundation, through grants IIS-0120663 and DMI-0522382, and the support of the Georgia Tech Rapid Prototyping and Manufacturing Institute member companies for sponsoring this work.REFERENCES[1] M. A. Mont and D. S. Hungerford, "Proximally Coated Ingrowth Prostheses: A Review," Clinicalorthopaedics and related research, vol. 344, pp.

A. Mont and D. S. Hungerford, "Proximally Coated Ingrowth Prostheses: A Review," Clinicalorthopaedics and related research, vol. 344, pp. 139-149, 1997.[2] "A Patient's Guideto Total Hip Replacement Surgery," Medical Multimedia Grouphttp://www.healthpages.org/AHP/LIBRARY/HLTHTOP/THR/INDEX.HTM, October 16, 2005 [3] "Hip Replacement," Biomet, Inc. http://www.biomet.com/patients/hip_replacement.cfm, October16, 2005[4] C. R. Bragdon, M. Jasty, M. Greene, H. E. Rubash, and W. H. Harris, "Biologic Fixation of TotalHip Implants," Journal of Bone & Joint Surgery, vol. 86, pp. 105-117, 2004.[5] "Code of Federal Food and Drug Administration Regulations: Hip joint metal/polymer/metalsemi-constrained porous-coated uncemented prosthesis," Wikipedia, Rockville MD, 2006.[6] M. Jasty, C. Bragdon, D. Burke, D. O’Connor, J. Lowenstein, and W. Harris, "In vivo skeletal responses to porous-surfaced implants subjected to small induced motions.," Journal of bone and joint surgery, vol.

Bragdon, D. Burke, D. O’Connor, J. Lowenstein, and W. Harris, "In vivo skeletal responses to porous-surfaced implants subjected to small induced motions.," Journal of bone and joint surgery, vol. 79, pp. 707-714, 1997.[7] S. Suresh, "Graded Materials for Resistance to Contact Deformation and Damage," Science, vol.292, pp. 2447-2451, 2001.[8] Nancy Elftman, "Orthotic Materials: Cellular urethanes combine durability with breathability," inBioMechanics, 1998.[9] M. Viceconti, "Visible Human Male - Bone surfaces," Laboratorio di Tecnologia Medica, Istituti Ortopedici Rizzoli http://www.tecno.ior.it/VRLAB/researchers/repository/BEL_repository.html, October 01, 2005[10] W. Sun, B. Starly, J. Nam, and A. Darling, "Bio-CAD modeling and its applications incomputer-aided tissue engineering," Computer-Aided Design, vol. 37, pp. 1097–1114, 2005.[11] A.

253, pp. 1182–1187, 2002.[18] H. Kita, M. Fukushima, D. D. Jayaseelan, Y.-P. Zeng, and K. Ohsumi, "Oxidation Resistance and High-Temperature Lubricating Properties of Magnesium-Phosphate-Treated Graphite," Journal of the American Ceramic Society, vol. 88, pp. 2632-2634, 2005.122[19] H. Wang and D. W. Rosen, "Computer-aided design methods for additive fabrication of truss structures," International Conference on Manufacturing Automation (ICMA2002), Hong Kong, China, pp. 181-198, December, 2002[20] P. Jacobs, Rapid prototyping & manufacturing: fundamentals of stereolithography, 1st ed,SME, Dearborn, MI, 1992.[21]J.-P. Kruth, P. Mercelis, J. V. Vaerenbergh, L. Froyen, and M. Rombouts, "Binding mechanisms in selective laser sintering and selective laser melting," Rapid Prototyping Journal, vol. 11, pp.

Kruth, P. Mercelis, J. V. Vaerenbergh, L. Froyen, and M. Rombouts, "Binding mechanisms in selective laser sintering and selective laser melting," Rapid Prototyping Journal, vol. 11, pp. 26-36, 2005.[22]"A NEW world-class generative process developed for direct product realisation with http://www.mcp-SELECTIVE group.com/rpt/rpttslm.html, October 15, 2005LASER MELTING(SLM)," MCPGroup[23] D. W. Hutmacher, "Scaffolds in tissue engineering bone and cartilage," Biomaterials, vol. 21,pp. 2529-2543, 2000.[24] D. W. Hutmacher, "Scaffold design and fabrication technologies for engineering tissues — state of the art and future perspectives," Journal of Biomaterials Science -- Polymer Edition, vol. 12, pp. 107-124, 2001.[25] H. Wang, "A unit cell approach for lightweight structure and compliant mechanism," PhD Dissertation, School of Mechanical Engineering, Georgia Institute Of Technology, Atlanta, GA, 2005.[26] H. Wang, Y. Chen, and D. W.

Wang, Y. Chen, and D. W. Rosen, "A Hybrid Geometric Modeling Method For Large Scale Conformal Cellular Structures," ASME Computers and Information in Engineering Conference, Long Beach, California, September 24-28, 2005, 2005, Paper Number: DETC02/CIE-34495[27] H. Yoshida, A. Faust, J. Wilckens, M. Kitagawa, J. Fetto, and E. Y.-S. Chao, "Three- dimensional dynamic hip contact area and pressure distribution during activities of daily living," Journal of Biomechanics, vol. 39, 2006.[28] M. Niinomi, "Fatigue performance and cyto-toxicity of low rigidity titanium alloy, Ti-29Nb-13Ta-4.6Zr," Biomaterials, vol. 24, pp. 2673-2683, 2003.123.

SIGN OF A HIGH TEMPERATURE rrnJ....,Ih;,~~for theSELECTIVE LASER SINTERINGJohn McWilliams, Christopher Hysinger,Department of Mechanical Engineering The University of Texas at AustinAustin,78712The quality of parts made by the Selective depends directly on controlling heat transfer to we detail the thermal design of a HIGH to be used for building parts from metals and r>A,Nll1t"l,." experimental techniques are used to .....""'-'".:;;.." transfer schemes to apply uniform flux to the completion of the project, we expect to have a test-bed to radiant and convective heat transfer on a environment.qualityBACKGROUNDBackground oftbe HTW ProjectIn 1989, work began on a new SLS machine to ceramics. This High Temperature Workstation (HTW) would to be heated to sintering temperatures in excess of temperatures is to use a high-power laser.

This High Temperature Workstation (HTW) would to be heated to sintering temperatures in excess of temperatures is to use a high-power laser. laser and to design a workstation around this high-power laser.decision wasIn polymer systems, experience has shown that powder-bed often causes the newly formed layer to deform so plane of the part-bed. This phenomenon is known as "curling". heat the part-bed has been found to control the curling problem1991. This workstation utilized the 1.1 kW CCh laser, but had no heating system to pre-heat the powder. Single layer experiments were done using stainless steel copper powders. Multiple layer experiments were then conducted using two alloy systems. [Wu] heating, no built at base of the conditions. This As suspected, curling was a problem.

Multiple layer experiments were then conducted using two alloy systems. [Wu] heating, no built at base of the conditions. This As suspected, curling was a problem. Since this chamberThe first High Temperature Workstation was brought onThe Next Stage: High Temperature Selective Laser SinteringThis paper details the design, development, and testing of a HIGH TEMPERATURE PROCESS CHAMBER, the key component of the new HTW presently construction. This phase of the HTW project required an evaluation of process requirements, the selection design, construction, and of a Heating Method and Temperature Control System, and Specifications for the testing of a HIGH TEMPERATURE PROCESS CHAMBER. First, HIGI-I TEMPERATURE PROCESS CHAMBER were developed. conceptual designs for a Heating Method were developed. The final conceptual design was evaluated and optimized using heat to accommodate the heating method.

conceptual designs for a Heating Method were developed. The final conceptual design was evaluated and optimized using heat to accommodate the heating method. Finally, the completed chamber was performance-tested to identify any potential problems, and to compare real performance to model predictions.transfer models. The PROCESS CHAMBER wasand built110DESIGN SPECIFICATIONSA Design Group was formed to develop Design Specifications for the HI G H TEMPERATURE PROCESS CHAMBER. Researchers in the Material Science area formed an important part of the design group, since they would be the end users of the HTW. The group developed three important Design Specifications.Maximum Temperature SpecificationAfter examining some of the properties of the materials to be used in the HTW, the Design Group set a goal for a Maximum Operating Temperature of 1300°C.

This value represents the final steady-state temperature that the powder-bed should reach with the heating system operating at full capacity. The group then set a Normal Operating Temperature range of 100°C to 10OO°C. We define Normal Operating Temperatures as those temperatures which fall within the controllable range of the heating system. Temperature Uniformity SpecificationA high degree of Temperature Uniformity is required to eliminate curling. The Temperature Uniformity Specification was therefore one of the most important design considerations, but also one of the most difficult to quantify. Since surface temperature uniformity depends on many factors and is material dependent, the group decided to evaluate the potential uniformity of each heating method based on its heat flux uniformity. Since there is no direct relationship between heat flux and temperature distribution, no value was set for this requirement.

Since there is no direct relationship between heat flux and temperature distribution, no value was set for this requirement. Rather, the Design Group made the decision to simulate the heat flux distribution for different heater designs, and to use the results as an important selection criterion when choosing between different heating methods. Controllability SpecificationThe controllability of the heating system depends on the dynamic properties of both the heating system and the control system. At this point in the design process, the design group could not quantify the Controllability Specification, but decided that it should be an important criteria in selecting a heating method.Heaters vary greatly in thermal mass. In low thermal mass systems, the power output can be changed rapidly. In high thermal mass systems, power output changes very slowly in response to power input.

In low thermal mass systems, the power output can be changed rapidly. In high thermal mass systems, power output changes very slowly in response to power input. A heating system with a low thermal mass is most desirable for flexibility in process control.HEATING METHOD SELECTION AND ANALYSIS In addition to the Design Specifictions given above, several other design constraints made selection of a heating method quite difficult. For example, the heating method had to allow the laser to scan the entire area of the part-bed. Simple models had shown that a very high heat flux is required to meet the Maximum Temperature Specification. In the selection process, the design group examined many high-power heating methods, including quartz infrared heaters, graphite ring heaters, direct resistance heating and induction heating. Selection and Application of Dual Infrared Panel HeatersThe design group selected a dual infrared panel heater arrangement as the heating method for the HTW (Fig. 1).

Selection and Application of Dual Infrared Panel HeatersThe design group selected a dual infrared panel heater arrangement as the heating method for the HTW (Fig. 1). A similar system was developed by Brubaker for an infrared drying oven. High-density quartz panel heaters were selected. The system has a total power output of around 20 kW.The design group chose the dual panel heater arrangement as the most promising design for several reasons. Most importantly, quartz heaters can deliver a higher surface heat flux than any other of the heating methods investigated (up to 780 Watts per square centimeter at the heater surface). Also, Brubaker's work suggested that the orientation of the heaters could be optimized to obtain a uniform heat flux. Finally, these heaters have a very low thermal mass which gives them excellent controllability.

Finally, these heaters have a very low thermal mass which gives them excellent controllability. Their low thermal mass allows them to reach full power output in about one second.111Thermal Modeling and Analysisfeasibility of the heating method could be evaluated. After we determined that the heaters could provide enough heat flux, two issues remained to be investigated. First, the heaters needed to be oriented to produce a relatively even heat flux across the part-bed. If it was possible to combination of obtain an even flux, then an orientation had to be chosen to obtain the heat flux uniformity and heater efficiency.

If it was possible to combination of obtain an even flux, then an orientation had to be chosen to obtain the heat flux uniformity and heater efficiency. The following model was to help achieve both of these goals.Now that commercially available heaters had been found,For radiant heat transfer between two surfaces A and Bspace,heatsurface A to surface B is directly proportional to Fa-b, the view factor from A to implies that if the view factors could be calculated, then the heat flux uniformity between the two surfaces could be evaluated. In our case, surface A and surface B is the powder-bed.A simulation program was developed that calculatesplanes, representing the radiant heaters, to an array of representing the powder-bed (Fig. 1). The two heaters are angled into powder-bed at an angle ft. The outside edge of the two heater planes is at a the powder-bed. The heaters are separated by a distance d.

1). The two heaters are angled into powder-bed at an angle ft. The outside edge of the two heater planes is at a the powder-bed. The heaters are separated by a distance d. The program plots view factor distribution across the plane of differential elements representing the powder-bed. The plane is the same size as the projected area of heaters. An efficiency is percentage of infrared radiation that strikes this plane.a uniform heat flux could be obtained for certain orientations. Figure 2 shows view-factor profile for the final chamber design. At the center of the bed, the view factor is approximately 0.40.

Figure 2 shows view-factor profile for the final chamber design. At the center of the bed, the view factor is approximately 0.40. For this configuration, the model predicts that 17% of the energy from the heaters should directly reach the powder-bed.After evaluating several different heater orientations, thepredicted2015>0 1051020 15 X Distanee(em)25Figure 1:Heater ConfigurationFigure 2: VeN Factor Contour Pbt112as~WINDO!W CYLINDERWINDOW CAPAIRINPROCESS GAS INWINDOW CYLINDERCylinder is to keep the laser has shown that the windowto be cooled. '"V.IH.......''U to accumulate deposits from inside up to the window by free the wall of the Window produced the new'l.l':ll·nn·..." wereDesi gn Sol utionAn aluminum cylinder and cap were constructed to hold the laser window (Fig. 4). The window holder has provisions for gas flow across both sides of the window.

4). The window holder has provisions for gas flow across both sides of the window. The inside surface of the window cylinder has a gap that allows process gas to flow across the entire exposed surface of the window.The cap also has provisions for gas flow, in this case across the outer surface of the window. Clean air is directed at the window through two jets, located at opposite ends of the cap.Gas Handling System Gas flow through the chamber and across the part bed will affect temperature distribution and part quality. An experimental approach was used to help design the Gas Flow System for the HIGH TEMPERATURE PROCESS CHAMBER.Design considerationsWe performed experiments with a thermal imager to determine the effect of convective flow on temperature uniformity across the powder-bed. The imaging experiments showed that free convection provided the most uniform temperature distribution across the powder-bed.

The imaging experiments showed that free convection provided the most uniform temperature distribution across the powder-bed. The procedure and more complete results are given in the System Performance section.Keeping in mind that gas flow must be introduced into the window cylinder to keep the window cool and clean, we used the experimental methods detailed below to help develop a gas flow system. The primary design consideration was to ensure that any gas flow scheme left the gas above the part-bed undisturbed so that it experienced only free convection. In the design of the Gas Flow System, another important consideration was to obtain efficient purging of oxygen and sintering by-products from the chamber.Experimental Row AnalysisIn order to visualize the flow patterns established inside the HIGH TEMPERATURE PROCESS CHAMBER, a mock-up of the chamber was constructed.

This test chamber was a I: I scale model of the actual chamber, but was constructed of clear polycarbonate sheet. The mock-up consisted of a window cylinder and the main chamber. In the actual chamber, the part bed would lie directly below the laser window cylinder in the plane of the plenum, which is an inch and a half above the base of the chamber. For a tracer, dry ice and water were placed in a cylinder located where the part bed would be. The boiling of the dry ice created C02 vapor. This vapor allowed us to observe the flow patterns over the part bed.At this point, the greatest concern was to determine how flow from the window cylinder would effect the bed. We developed and tested a simple flow scheme (Fig. 5). In this scheme, all of the gas flow into the chamber entered through the window cylinder. The gas was drawn around the part-bed through holes in each plenum, which were located in a symmetrical pattern to either side of the part cylinder.

The gas was drawn around the part-bed through holes in each plenum, which were located in a symmetrical pattern to either side of the part cylinder. The gas then exited the chamber through an exhaust port in the feed-side end of the chamber.Design SolutionWe found that for gas flow rates of less than 40 liters per minute, a stable flow pattern could be achieved around the part bed with the simple flow scheme described above. At these flow rates, we observed that the flow patterns of the CO2 vapor over the part-bed remained undisturbed. From this observation, we concluded that the gas around the part bed also remained undisturbed, which was the primary design goal for the Row System.In the HI GH TEMPERATURE PROCESS CHAMBER, we plan to install two cross flow heat exchangers, one on each side of the part-bed.

In the final system, gas will enter the chamber through the Window Cylinder, pass through the fins of the heat exchangers, and travel beneath the plenum to the feed-side of the chamber, where it will exit through an exhaust port. The purpose of the heat exchangers is to absorb the excess heat from the chamber and to cool the gas before it is released to the building exhaust system.114SYSTEM PERFORMANCE When the chamber was completed, we developed a test plan to evaluate the performance of the chamber before it was installed in the workstation. Performance objectives were first developed, and then a series of tests were designed to evaluate the performance of the HIGH TEMPERATURE PROCESS CHAMBER.Extended Burn-in TestThe main purpose of the Extended Burn-in Test was to determine if the heater windows and seals could withstand the temperatures to which they would be subjected.

Another purpose of the test was to examine the affect of insulation on chamber and seal temperatures.In this simple test, both ends of the chamber were removed, and a table-fan was used to blow air through the chamber for convective cooling. The heaters were run for 30 minutes, and then the condition of the window seals was inspected.The first time this test was run, the Viton1'M o-ring seals started to burn after about 10 minutes of heater operation at full power. The second time the test was run, the ceramic fiber insulation was installed inside the chamber. This time, the seals remained intact, and the chamber walls remained much cooler. This indicates that the seals absorbed most of the heat from the chamber walls, not directly from the heaters. Water Bath TestsThe main purpose of the Water-bath Tests was to determine the efficiency of the heating system.

Water Bath TestsThe main purpose of the Water-bath Tests was to determine the efficiency of the heating system. In this case, we have defined efficiency as the proportion of the energy supplied to the heaters that reaches the heated surfaces.Two tests were performed. The objective of the first test was to determine the heat flux across the entire area of the part-side of the chamber module base. In the second test, the objective was to determine the heat flux across the part-side of the powder-bed only. For each test, a tray containing a known weight of water was placed in the chamber, the heaters were turned on for 30 seconds, and the change in temperature was measured.In the first water-bath test, the chamber base received an average of 134 kJ of energy. When averaged over the 30 seconds of the test, this works out to an average power 4.47 kW. The heat flux per unit area was 41 W/in2 or 6.4 W/cm2. The electric power input to the heater was maintained at 16.6kW.

The heat flux per unit area was 41 W/in2 or 6.4 W/cm2. The electric power input to the heater was maintained at 16.6kW. This yields an efficiency of 27%, which means that 27% of the electrical energy input to the heaters was converted into heat on the part-side of the chamber base.The results of the second test indicate that the powder-bed will receive an averageheat flux of 8.58 W/cm2 , yielding an efficiency of 18% on the powder-bed.IR Scanner TestsOne of the most important design criteria for the heating system was to produce very uniform bed temperatures. In measuring the temperature distribution, we hoped to accomplish two goals. First, we wanted to see if the measured distribution coincided with what we expected from the View Factor Optimization. The second goal was to see how different parameters affected temperature uniformity.

The second goal was to see how different parameters affected temperature uniformity. Tests were performed to determine how temperature distribution was affected by gas flow across the powder-bed and by chamber wall reflectivity.To perform the IR scanner tests, the IR scanner was placed on top of the chamber, where it looked down the hole where the window cylinder would normally be installed. During each test, the heater power was kept constant by maintaining heater power at 30 A. A 3-5 Jim filter was used to minimize the effect of reflected IR energy on the temperature reading. In each test, the specimen and chamber were set up and the heater power was turned on. After the temperature profile reached steady-state, the output from the IR scanner was recorded on video tape.In one of the most important tests, the effect of convection across the powder-bed was investigated. In this test, the flow velocity across the powder-bed was varied using a116InfraredSystem..

sign of a Higb Temp.erature Workstation for the Selective Laser Sintering ProcessSuman Das, Johll. McWlllia.tns, BennyV¥u, J. J. BeatnanDeparttnentofMechanical Engineering The University ofTexas at Austin Austin TX 78712AbstractThe Selective Laser Sintenng ProcesshassuccessfuUy proven its viability for •sinteringlow meltin~·iPointmeta1s,polymers, and.ceramic binder mix tures. A High'femperatureWorkstation utilizing a high-power laser.is cur rently beingdesignedtQstudythefeasibilityof this process for high melting-point>lllaterials.Aworkingqhamberand powder handlin~ system were designedtoacceptpowerfrollla 1.1 kW laser. A powder preheating systelll.andgas·flOW·COlltr01.systemregulate. the chamber.environment. A process control computeraut01l1ates all aspects of the ·High Temperature Workstation. At the cOlllpletiQnoftheproject, we expect to have an auto mated academic test-bed to control and monitor all process parameters associated.

At the cOlllpletiQnoftheproject, we expect to have an auto mated academic test-bed to control and monitor all process parameters associated. with high-temperature Selective LaserSintering.BackgroundIn 1989, the latest in a series of •• Selective Laser Sinteringmachineswas designed and built at The.University of 'fexasatAustin. [1] The machine uses a low power laser (lessthanlOOW) to build Parts frollliowmeitingpoint materials such as polymers, waxes, low-me1ting-pointmetals,andceratn1c-pinder.mixtures. It allows minimal control of the sinteringenvironment, and operates at near atmospheric pressures and temperatures. A smallradiant heater is used to preheat the powder bed.Desigll.work.has. begunOll..anew•High·Temperature Workstation (HTW). The new workstation wiJ1use a high-.pQwerC021aser to sinter such materials as steel and ceramics.

begunOll..anew•High·Temperature Workstation (HTW). The new workstation wiJ1use a high-.pQwerC021aser to sinter such materials as steel and ceramics. It wilLalso feature a much lllQre flexible environmental control system, incorporating pressure, gas flow, and temperature control systems.Design ObjectivesHi&hJ.,aaer powerUntilnow,the Selective Laser Sintenng project at the University has focussed on low I1leltingpointlllatenals,such as polymers and low melting point metals. These materials are useful fOf1l1akingprototypeparts, .or for. making preforms·· of parts which require. post processing.OneQf the goals ofthe HTW project is 10 directly sinter structural parts which have materiaLproperties cOI1lparable to thQse producec1 using conventional maqhinin~operations. This implies havingithe capability·to·directly sinter materials such as steel andCo-Wcermet. To sintersteel,thepowderlllustreach temperatures over 1200°C.

This implies havingithe capability·to·directly sinter materials such as steel andCo-Wcermet. To sintersteel,thepowderlllustreach temperatures over 1200°C. One way to reach these teIl1peratures is to use a high-power laser.Another active area of SFF research involves the direct sintering of ceramic parts. Presently, Selective Laser Sintering is being used to make ceramic preforms. The preforms consist of a ceramic powder, such as alumina, which is bound together by a polymer164binder. The preform is post-processed by firing it in an oven. Using a high-powered laser, it may be ossible to directly sinter ceramics without using a binder or post-processing.benefit of using high laser power is to achieve a maximum scanning rate. A maximum scanning rate implies that the scanning speed is limited only by the scanning mechanism and not by laser power. A maximum scanning rate is desirable for two reasons.

A maximum scanning rate implies that the scanning speed is limited only by the scanning mechanism and not by laser power. A maximum scanning rate is desirable for two reasons. First, it decreases process time to help achieve to goal of true rapid prototyping. Second, higher scanning speeds generally produce better quality parts due to improved heat transfer characteristics.AHiah Temperature EnyironmentThe major goal behind the HTW project is to directly sinter high-melting-point materials. One way to do this is to use a high-power laser to provide the heat necessary to melt the powder. Another way is to preheat the powder until it is close to its melting point. A low-power laser may then be used to supply the heat of fusion necessary to sinter the pre-heated powder.The High Temperature Workstation will incorporate a heating system to pre-heat the powder bed to a maximum temperature of 1300°C, with a controllable range of 100°C to lOOOoC.

Even with a high-power laser, pre-heating is desirable for several reasons. Preheating is currently used to eliminate thermal stresses which are brought. about by uneven cooling after each layer is sintered. These stresses can lead to deformation of the part known as curling. The curling problem is likely to be more severe when high temperature materials are sintered, since higher temperature gradients will be present. powder bed also will allow the laser to scan the part more rapidly when Preheating sintering . -temperature materials, since the laser energy can be used primarily to melt the preheated powder. Finally, for very high-temperature materials such as ceramics, the high-power laser alone may still not provide enough heat to sinter the material. Pre-heating may be necessary if these materials are to be directly sintered.Gas Flow ControlThe working chamber of the High Temperature Workstation will incorporate a sophisticated atmospheric control system.

The most important function of the atmospheric control system to maintain a safe atmosphere in the chamber. The role of gas flow control in the overall safety design of the HTW will be addressed in the Safet;x section. A gas flow control system is also important for material processing research. The final important function of the gas flow control system is to remove the by-products of sintering from the chamber.From a materials processing standpoint, an atmospheric control system will enable the study of the effects of different gaseous environments on part quality. Also, it allows the possibility of layer-wise gas-phase material processing, such as carburizing or nitriding a steel part as it is sintered.Finally, the atmospheric control system must maintain a relatively clean environment inside the chamber.

That is, a sufficient quantity of gas must pass through the chamber to keep it relatively free of sintering byproducts, which will tend to foul the laser window and other internal components of the chamber. Also, a hazy environment in the chamber will reduce the amount of laser energy which reaches the part, and will interfere with observation of the process through the viewing window.Vacuum EnvironmentA future goal for the HTW is to have a high-temperature, high-vacuum sintering chamber. A vacuum environment will eliminate oxidation problems, and may improve part densities. Most conventional sintering is donevacuum sintering ovens.165PrQcess CQntrQIOne Qf the design Qbjectives at the start Qf the Selective Laser Sintering prQject was tQ design a fully autQmated machine which CQuid be used fQr rapid prQtQtyping. Unlike cQnventiQnal machining QperatiQns, parts made by Selective Laser S~ntering were tQ require a minimum of human intervention and expertise.

Unlike cQnventiQnal machining QperatiQns, parts made by Selective Laser S~ntering were tQ require a minimum of human intervention and expertise. To approach this goal, the HTW will incorporate a sophisticated Process CQntrol System. At the heart Qf this system, the prQcess cQntrol computer will autQmate all aspects Qf machine operation, including laser and scanning system, mechanical actuators which drive the pQwder feed and pQwder leveling mechanisms, and gas flow cQntrQI. The PrQcess CQntrQI CQmputer will alsQ perform data logging Qf all prQcess parameters. In the future, the PrQcess CQntrQI System will mQnitQr and cQntrQI prQcess temPeratures and pressures.~The Selective Laser Sintering PrQcess has had a very gQQd safety recQrd tQ date, but dQes have the pQtential fQr safety problems if SQme simple safety precautiQns are nQt followed. All of the pQtential hazards associated with the 10w-pQwer sintering machine will be present in the HTW.

All of the pQtential hazards associated with the 10w-pQwer sintering machine will be present in the HTW. In additiQn, the high laser pQwer and greater variety Qf materials and forming gases used create additional hazards which much be cQntrQlled.Many types of materials are pQtentially explQsive in pQwder fQrm. According tQ the American SQciety fQr Testing and Materials standard,in rare circumstances, there is a pQtential for dust explQsiQn in the HTW. AccQrding tQ Dr. K. Chatrathi [2], simply speaking, three elements are required to create a dust explQsion: dust and an oxidizer in certain prQpQrtiQns tQ each Qther, and a strQng ignitiQn SQurce. Our strategy will be tQ eliminate the Qxidizer frQm the chamber. With the high activatiQn energy supplied by the laser, it will be necessary to maintain an inert environment in the chamber to eliminate the chance of an explosion.

With the high activatiQn energy supplied by the laser, it will be necessary to maintain an inert environment in the chamber to eliminate the chance of an explosion. A gas delivery system, and a gas sensing system equipped with an alarm will be necessary to minimize any chance for an explosiQn.The high-pQwer laser tQ be used in the HTW is extremely hazardous to bQth the eyes and the skin. The best way tQ minimize the laser hazard is tQ maintain a cQmpletely enclosed system that will nQt allow any laser radiatiQn tQ escape the machine under normal operating conditiQns. GQvernment regulatiQns also specify that the laser system is tQ be tQtally enclQsed if at all pQssible.

GQvernment regulatiQns also specify that the laser system is tQ be tQtally enclQsed if at all pQssible. A safety interlQck system will ensure that nQ laser radiatiQn can escape frQm the HTW under nQrmal QPerating cQnditiQns.Design StrategyIn order to evaluate the feasibility Qf High Temperature Selective Laser Sintering, the design was divided intQ three phases.Phase I: Hi&h Laser PQWer SinterinaIn Phase I, the feasibility Qf high-laser-pQwer sintering will be evaluated. At the completion of this phase of the project, the high-power laser, and its associated scanning and control systems, will be operational. A gas tight chamber complete with powder handling system will be tested. Sintering will be carried out in inert environment at atmospheric temPeratures and pressures.PhaSf( ll: Wih Temperature SinteriniIn Phase ll, the powder surface will be preheated to high temperatures.

This phase involves the design Qf a powder heating system and its associated temperature measurement and control systems. Phase II will likely require a redesign of the chamber, since it will have to accommodate the new heating system, and will have to be designed with thermal performance in mind The powder leveling system will also have to be redesigned to handle the high temperatures.166Phase ill: vacuum SinteriniIn phase ill, sintering will be carried out in a very low pressure environment, as it is in conventional vacuum sintering processes. The high temperature capability of Phase IT will be maintained. This final phase will require the design of a third chamber, one which can handle both vacuum and high temperature.Design FeaturesAt this time, the design for the fIrSt phase of the H1W has been completed, and fabrication is well underway.

The following sections outline how the designers met the Design Objectives previously discussed for Phase 1of the project.Hiih-Power CO2 LaserHigh-melting-point materials, such as steel and alumina powders, will be primarily used in the High Temperature Workstation. Thus a high-power laser is required. The reasons for selecting a 1100 Watt CO2laser (wavelength 10.6 J.ltll) are:1. One of the highest efficiencies for the conversion of electrical to optical energy is found in the CO2 laser. This property, together with ease of operation at high power levels, has made the CO2 laser widely applicable as a source in the laser processing of materials.2. The cost of a C02 laser is less than for any other laser at the same power level. 3. C02 lasers are presently used to sinter powders in laser cladding. 4. Most ceramic materials have good absorption of CO2 laser energy.

3. C02 lasers are presently used to sinter powders in laser cladding. 4. Most ceramic materials have good absorption of CO2 laser energy. Some materials with Si-O bonds are relatively transparent in the visible but absorb strongly near 10 J.1m.For metal materials, theoretically the absorption of laser energy by metal surfaces at 20°C is much lower at infrared wavelengths than at visible wavelengths. However the low absorption values apply only for clean metal surfaces. With the trapping of powder surfaces and at high-temperature conditions, the values can be increased substantially. In addition, from W.W. Duley[3] , Laser Processing and Analysis of Materials,"many experiments show that a substantial increase in the coupling coefficient (equivalent to absorption mte) for laser mdiation at the target occurs when the intensity is sufficient.

This p a s some fraction of to initiate a breakdown plasma in the local the incident laser radiation and transfe to the target via hydrodynamic expansion.

While the energy coupled from energy to the target can be increased by this mechanism, it is deposited over a large area than the normal beam focus."the resulP£AK INTlN51N IW ".-2,From figure 2.8 [3], we see a dramatic increase of coupling coefficient and decrease ofreflectivity at laser the 1 Watt C02 laser at the peak power output was calculated using the condition: Laser peak power output 5500 Watts, The radius of the minimum spot size rs is 0.003 cm.intensity 107 W/cm2·The peak intensity of2P 10=-= 1ter22x5500s 1txO.00328=3.89xI0 W/cm 2[3]i! ~ ~ 01 i1008060401O10'10110·415' IIllS 5'111Reflectivity 01...._C~'"'l C_dlnt.i., /0 11l" { ~z.JJjJThis result shows that the C02 laser has sufficient energy to create a plasma which dramatically changes metal reflectivity and increases coupling rate.10100ENERGY OEN5'lY <Jem -2,."1..,. La. The redectivily and thermal coupling coefficient o( AISI 1045 sleel at 10.6 ~m.

La. The redectivily and thermal coupling coefficient o( AISI 1045 sleel at 10.6 ~m. The Jl(ak IPU!lnJiIY UI ellamated on the basislJf a O.5·,U$ pulse duration IRoeulcr an<l Gregson (1978)).167Gas Flow SystemIn the initial phase of the project, the sintering process will be carried out at atmospheric pressures. An inert gas, either nitrogen or argon, will be fed into the chamber through the gas flow system. A residual gas analyzer will continuously sample the process gas in the chamber. Oxygen content will be monitored for safety purposes. The gas analyzer is tied into an alarm system which will shut down the process if the oxygen partial pressure exceeds a pre-determined limit.Mass flow controllers will control the proportions of up to four different process gases. An inert carrier gas will always be used. In addition, up to three other process gases may be fed into the chamber through separate mass flow controllers.

An inert carrier gas will always be used. In addition, up to three other process gases may be fed into the chamber through separate mass flow controllers. These gas flows will be sensed by the gas analyzer, which will send a control signal to the mass flow controllers. The gas flow system is compatible with corrosive gases, such as ammonia. With additional hardware, the gas flow system can be adapted for use in a vacuum sintering environment.Safety Interlock SystemIn the design of safety systems for the High Temperature Workstation, safety interlocks are required for both the laser and for the chamber atmospheric conditions. Both interlock systems are coordinated by the data-logger, which also functions as an alarm annunciator and safety interlock relay system. With this system, all safety interlocks function independently of the process control computer.

With this system, all safety interlocks function independently of the process control computer. When a safety interlock is tripped, the data-logger sends an analog alarm signal which activates a solid state relay connected to the shutter control on the laser, effectively containing the laser energy without having to shut down the laser. All interlocks are fail-safe, and incorporate a manual reset, so that the process cannot resume until the alarm is in the safe state and the interlock has been manually reset. Manual overrides are provided for equipment calibration and testing.The first and perhaps most critical safety issue is the containment of laser radiation from the CO2 laser. Under normal operating conditions, no laser radiation can be permitted to escape the machine. To prevent accidental exposure, the chamber door is safety interlocked. The door which provides access to· the laser's optics is also interlocked.

To prevent accidental exposure, the chamber door is safety interlocked. The door which provides access to· the laser's optics is also interlocked. Magnetic switches are used on both doors.The second area of concern is the explosion hazard associated with powdered materials exposed to a high-energy source such as a laser or heater. Under normal operating procedures, all sintering will be carried out in an inert environment. Should oxygen inadvertently leak into the chamber, it will be detected by the gas analyzer, which constantly monitors oxygen concentration and sends an alarm signal to the data-logger when a pre-set concentration limit has been exceeded.Laser Scannini SystemThe scanning system is driven by a set of X-Y linear motion tables with a positioning accuracy .of 0.001 tf over a travel of 12 ".

These tables move the mirrors that direct the laser beam onto a moving focus lens assembly over the area to be irradiated.The motion of the tables is controlled by a high precision CNC controller. This system scans the laser orthogonally over the work area unlike the conventional mirror scanning system which requires a complex correction for its lack of orthographic projection. Our application requires a constant speed of scanning and laser scan speeds of up to 20 inches/sec will be attained by this system. The total scan area of this system will be 12" X 12".168Proces§ •Control S·Y§~m The design ofthe High temperature workstation requiresc.omputer con-· trol of a number ofdevices.•• Amortg the devices to be interfaced to the process control computer are:O.C.motor Control"'rRoller Actuatora. Stepper motor andD.C motor controllers for the· powder deliv ery and leveling system.Power supply __--jb. A CNC· controller for the laserscanner.c.

Stepper motor andD.C motor controllers for the· powder deliv ery and leveling system.Power supply __--jb. A CNC· controller for the laserscanner.c. Adatalog~er which reports variFrom 6mk switchesoustemperatDtes.d. A chatnt>etgasanalyzer. e. A safety·interlock for the process-RS232CserialOata LoggerInpulsing cha.rober.RS232Cll8rialf. Temperature and heatercontroller.RS 2320 serialThe aim of the process control software is to implement an integrated process control system which automates all aspects. of the SLSprocess and alloW's the user to set process parameters, receive graphicalprocess.feedback.

of the SLSprocess and alloW's the user to set process parameters, receive graphicalprocess.feedback. and simulation.· of the sinteringprocess on screen.RS232Cll8rialFrom door swiloh§af$ly InterlockA schel1lJ1tic ofthe H'1W Process Control SystemLASER WINDOWPARTMODULECHAMBERUNEAR ACTUATORLINEAR ACTUATORSSchematic drawing of the High Temperature Workstation169Chamber DesimThe chamber design is based on ease and safety of operation, and is designed to provide environmental control to facilitate sintering. A gas flow pattern which does not affect the uniform temPerature distribution was considered. To distribute the gas flow uniformly across the powder bed, a sPecial ring distributor with small holes on its periphery was designed. An experiment showed that gas was evenly distributed 2 inches above the bed.

An experiment showed that gas was evenly distributed 2 inches above the bed. The gas exits from 4 ports placed at comers of the chamber, which ensure that the gas leaves the test bed evenly .A special laser window is required to allow the laser beam to enter the chamber, while maintaining a sealed environment inside the chamber. Several materials were investigated, including sodium chloride and zinc selenide. Zinc selenide was selected as the window material, because it possesses a high transmittance and durability and is the best overall material for a C02 laser. In order to prevent high-temperature radiation from the powder bed from damaging the window, the interior side of the window has a reflective coating which reflects a wide range of wavelengths.Project Status Currently, the chamber, powder handling system, and safety interlock system designs are complete. The laser and scanning system are to be delivered in August 1991.

The laser and scanning system are to be delivered in August 1991. Phase I of the system should be running in the Fall of 1991.Future WorkPhase D: Hi&h Tenwerature Sinterin&The next phase of the High TemPerature Workstation design will be to incorporate powder preheat capability into the chamber design. This will involve utilizing one or more heating methods to pre-heat the powder bed to a maximum temperature of 1000°C. The thermal design of the chamber and associated hardware is now in the conceptual design stage. Heating will likely be done be a radiant heater, or an internal heat generation scheme, such as induction heating. The method used will be equally effective at atmospheric or at vacuum conditions. The present Process Control System is already equipped to handle temperature measurement and control.

The method used will be equally effective at atmospheric or at vacuum conditions. The present Process Control System is already equipped to handle temperature measurement and control. This high-temperature phase of the design will likely be implemented by the beginning of 1992.Phase ill: Vacuum Sinterin&In the final design phase of the High TemPerature Workstation, total environmental control of the sintering process will be possible. The main feature of this final phase of the High TemPerature Workstation Project will be the capability to sinter in a medium to high vacuum environment (10-3 to 10-6 torr). Difficulties to be overcome include comPensating for the outgassing of the powder, and the problems associated with evacuating sintering by-products from the chamber. The potential benefit to vacuum sintering is the realization of high density sintered parts.References[1] Forderhase, Paul A. Design ofa Selective Laser Sintering Machine IntendedforAcademic Use.

Design ofa Selective Laser Sintering Machine IntendedforAcademic Use. Master's Thesis, The University of Texas at Austin, May 1989.[2] Chatrathi, K., "How to safely handle explosive dust-Part 1", Powder and BulkEngineering, January, 1991, p. 22-23.[3] Duley, W.W., Laser Processing and Analysis ofMaterials , Plenum Press, NewYork, NY, 1983, p. 74-75.170.

SIGN OF A MICRO-HOPPER ARRAY FOR MULTI-MATERIAL POWDER DEPOSITIONSuman Das and James Santosa Department of Mechanical Engineering The University of Michigan Ann Arbor, MI 48109-2125AbstractWe present a concept for creating patterned beds of multi-material powder particles using an array of small-scale hoppers. Using this technique, we propose that to place fine particles of multiple fluidized powders discretely in a thin layer as opposed to depositing an entire powder layer of uniform composition using a roller device or doctor blade. Processing and consolidation of multiple, patterned powders can enable fabrication of composite objects with spatially varying structural and multifunctional characteristics. Although theory on the design of small-scale hoppers is lacking, our design for a hopper, its valving, and its particle delivery system are guided by background theory for large hoppers.

Although theory on the design of small-scale hoppers is lacking, our design for a hopper, its valving, and its particle delivery system are guided by background theory for large hoppers. A hopper array configuration is proposed, and a calculation for deposition time is presented. Delivery of powder was achieved on a prototype hopper. Experimentally measured mass flow rates were used to justify the use of this hopper with SLS and to guide further design improvements.INTRODUCTIONThe variety of applications for multi-material composite structures is both well established and growing rapidly. However, a present obstacle to continued growth lies in the approaches to manufacturing these types of artifacts. The development of new approaches for fabricating multi-material components is in part dependent upon developing SFF techniques capable of depositing and consolidating multiple materials. In particular, the selective laser sintering (SLS) process is well suited to this challenge.

In particular, the selective laser sintering (SLS) process is well suited to this challenge. A means of incorporating the delivery of multiple powders in SLS may enable the production of complex, multi-functional and intelligent components.Figure 1. (a) Navajo sand painting, The Arizona State Museum, University of Arizona and (b) Conceptual design of a multiple material, patterned, powder layer for building multifunctional components by SLS.155The concept of 2-D deposition for multiple powders is by no means a new idea. In fact the greatest model for this idea is based on the art form of sand painting as shown in Figure 1a and first described in the context of SFF by Pegna et al [7]. The Navajo and Tibetan monks of earlier centuries have shown that detailed patterning of fine powders can be achieved manually. These artists typically spent several weeks working patiently on a single-layer design.

These artists typically spent several weeks working patiently on a single-layer design. However, the challenge from a layered manufacturing perspective is in successively depositing thin layers of multiple powders quickly and with high placement accuracy. By allowing the consolidation of multi-materials as shown in Figure 1b, SLS has the potential to launch a new manufacturing paradigm that enables fabrication of complex, multifunctional, intelligent components that cannot be fabricated by existing methods. The following sections provide background on the use of hoppers with fine powders, the theory guiding the design of a single hopper and standpipe to an array of these devices, and results of preliminary experiments conducted on a prototype hopper design.BACKGROUNDIn current selective laser sintering machines, powder is leveled over the part-build area by a counter-rotating roller mechanism or a doctor blade.

A bed of powder may be quickly rolled over the build area in layers of consistent thickness. Currently the SLS process is limited to However, SLS is capable of fabricating components of only a single composition. simultaneously sintering a variety of powdered materials (metals, ceramics, polymers). This makes depositing and processing two or more different powdered materials side-by-side on the bed feasible. With this in mind, the development of hoppers for delivering dots, sharp lines and prescribed patterns of powder is desirable for use in the SLS process. In fact, those who practiced the art of sand painting used a tool that is essentially a hopper releasing powder to a surface by vibration. A multi-powder deposition device for adaptation to SLS should have more stringent requirements. For the design of our hopper, a number of design specifications were established. According to these specifications, the device should be able to:1.

For the design of our hopper, a number of design specifications were established. According to these specifications, the device should be able to:1. Deposit arbitrary patterns of multiple powder material (i.e. points, lines, regions) for particles with diameters ranging from 0.1 to 100 (cid:181)m. 2. Achieve a minimum in-plane feature size of 100 (cid:181)m. 3. Achieve layer thickness ranging from 25 to 250 (cid:181)m. 4. Regulate hopper mass flow rate, start and stop flow with minimal time delay. 5. Handle multiple materials with real time composition control.In an attempt to meet some of these goals, the background research focuses on 3 areas: the relevant properties of powder particles, the design of hoppers, and, flow control and valving techniques.Relevant Powder PropertiesThe characteristics of powdered material that must be considered are the size, shape (sphericity), cohesion properties, bulk density, and porosity (voidage).

These are most critical to understanding the bulk flow of the solid material, from how it is packed to how it is transported (Shamlou [8], Woodcock and Mason [11]). Cohesive properties can be described as the ability of the bulk material to freely flow based on inter-particle forces (Jenike [3]). One of several methods to measure cohesion property that is of special concern to SLS is the angle of repose156(a.k.a. angle of internal friction) of a deposited powder. This is useful for understanding the ability of a poured powder to remain in a stable pile (Stepanoff [10]) i.e. when the deposited lines are settling on the substrate.For applications over a wide variety of materials, it is important to have a practical method for classifying powders without depending on extensive experiments. The Geldart classification [2] is an appropriate and well-used system for grouping powders by their ability to fluidize.

The Geldart classification [2] is an appropriate and well-used system for grouping powders by their ability to fluidize. Fluidization is the ability of a uniform bed of particles to expand when an upward stream of gases is distributed from beneath. A bulk material that is well fluidized will have an expanded volume, where particles seem to be floating in the fluid stream, separate from other powder particles. This is advantageous because of reduced inter-particle friction and ability to circulate powders within the bed.SLS applications require fine powder particle diameter between 0.1 (cid:181)m to 100 (cid:181)m, and the bulk material is largely considered non-cohesive (fair to free-flowing in nature). These types of powders would typically fall into the Geldart A classification [2]. An attempt to fluidize such powders relies upon the knowledge of minimum fluidization velocity.

These types of powders would typically fall into the Geldart A classification [2]. An attempt to fluidize such powders relies upon the knowledge of minimum fluidization velocity. This is the velocity of introduced gas at which incipient fluidization occurs in the powder bed or chamber. There are many approaches to gaining values for minimum fluidization velocity, but equation (1) is the rule of thumb for particles in air (Woodcock and Mason [11]).Umf≅420ρ pd2 v(1)where Umf = minimum fluidization velocity (m/s)ρB= bulk density of the powder (g/m3) dv = mean particle diameter based on volume (m) and 420 is an empirical value for particles in air (m2/g-s).Hopper DesignA hopper can be described as the lower converging section of any vertical bin for discharging bulk materials. Hoppers can have conical, wedge-shaped, or even asymmetrical designs. Hoppers can also have circular, square, or rectangular orifices.

Hoppers can have conical, wedge-shaped, or even asymmetrical designs. Hoppers can also have circular, square, or rectangular orifices. Evidently, the geometry of a hopper design is essential in achieving desired flow properties of the discharged material. The delivery of particles may happen under core flow (first in-last out) or the more desirable mass flow (first in-first out). For the design of a conical hopper, the hopper half-angle, α, and the orifice diameter, D0 are the most important dimensions to be considered in preventing arching (bridging of particles that blocks flow) and rat-holing (an empty tunnel over the orifice due to core-flow). Jing and Li [5] provide useful information on hoppers for fine powders, stating an effective hopper half-angle of 10.2(cid:176) or less for maintaining mass flow.One of the most frequently used theories on flow from hoppers is W. A. Beverloo(cid:146)s modified correlation [1] for mass flow rate in conical hoppers.

Beverloo(cid:146)s correlation states that as particle size decreases, there is an increase in the mass flow rate. Empirical results agree for particle sizes down to 500(cid:181)m, where actual mass flow rate takes on much lower values than predicted.An article by Spink and Nedderman [9] presents an experimental and theoretical study on the rate of discharge of fine particles (diameters between 50 (cid:181)m and 500 (cid:181)m) under gravity in air from a hopper. These size particles are small enough to be retarded by air resistance when passing through the hopper orifice, but not too small so as to be subject to cohesive effects. In this work, it is believed that the air pressure gradient opposed to the downward flow reduces the flow rate of the particles under 500 (cid:181)m, and it is also presumed that change in voidage as the hopper converges has a slowing effect as well.

The work by Spink and Nedderman used first principles (stress equations for different regions in the hopper) to come up with a theoretical model. They experimentally measured the voidage and interstitial fluid pressure distribution with external probes around the hopper and used these to account for the drop. The result was a downward sloping set of data points for the mass flow rate of particles having less than 500 (cid:181)m diameter (always higher than actual, but generally following the reducing trend). This is a good basis for the kind of experimentation needed to understand the behavior (voidage changes and pressure distribution) in much smaller hoppers. Our challenge in designing a micro-hopper based powder deposition system for SLS will be to adapt these types of experiments to small scale hoppers.At the present, most theoretical predictions relate to large, coarse particles on the order of 500 (cid:181)m diameter that are cohesionless and insensitive to drag.

Beverloo(cid:146)s equation is known to be inaccurate for particles under 500 (cid:181)m diameter. Most hopper theory is valid for large-scale bulk materials handling only. However, the existing theory provides a useful basis on which empirical models, and eventually a working theory for powders under 500(cid:181)m and orifices around 100(cid:181)m can be developed.Flow Control and ValvingThere are a number of techniques to starting/stopping flow, from the use of physical barriers such as gate valves, to fitting a hopper with flow aids like vibrating plates. Considering the small particle and hopper sizes we are compelled to modify or combine existing techniques for application to an array of hoppers.158Pegna et al [7] conducted an experiment using a non-mechanical L-valve [6], where powders come to rest in the elbow portion due to their angle of repose. Flow was started by introducing air from an aeration point.

Flow was started by introducing air from an aeration point. The flow rate was proven to be unstable considering the size of the particles. However, this design did not feature a hopper, and the orifice was simply the end of the downcomer. The concept of a hopper valve for Geldart A powders is presented in an article by Jing and Li [4]. In this setup, the flow regime and flow stopping are controlled by maintaining a negative pressure gradient from the hopper orifice to the standpipe. It is our belief that combining the design of a standpipe with L-valve and hopper dimensions according to the Jing and Li criteria can be a strong combination for use with a greater variety of powders.EXPERIMENTSFigure 2 shows our prototype hopper design. Some of the featuresinclude:1. A 60mm I.D. glass Buchner funnel and coarse fritted glass disk used as a chamber for fluidizing particles and delivering them to the standpipe.2.

Some of the featuresinclude:1. A 60mm I.D. glass Buchner funnel and coarse fritted glass disk used as a chamber for fluidizing particles and delivering them to the standpipe.2. Two angled aeration points at the base of the funnel to fluidize the powder above the fritted glass disk. Fluidized powder travels down the standpipe originating at a hole in the fritted disk.3. Adjustable tube lengths for the glass standpipe and elbow in 10 mm increments. Glass pieces are joined using photosensitive UV curing resin. The prototype shown has a 150mm long standpipe and 40mm long elbow, both 4mm I.D.4. An L-valve aeration point angled at 45” to the elbow. The position of the aeration point can also be varied along the standpipe with the use of adjustable tube lengths.5. Bio Plas, Inc. pipette tips used as micro-hoppers, which feature a half angle < 10.2° and 5mm orifice that can also be trimmed for larger diameters (.75 and 1mm).6.

Bio Plas, Inc. pipette tips used as micro-hoppers, which feature a half angle < 10.2° and 5mm orifice that can also be trimmed for larger diameters (.75 and 1mm).6. 100 (cid:181)m glass beads used for powder for initial testing (bulk density 2480000g/m3),120 grit sandpaper used as the substrate (depositing surface).Figure 2. Prototype micro-hopperRESULTSTen trial runs for mass flow rate were conducted on the prototype hopper under both gravity flow and air assisted flow for orifice diameters of 0.5, 0.75, and 1 mm. Our preliminary trials show that the standpipe with L-valve design is fully capable of stopping the flow of powders. The powder will come to rest at the angle of repose (approximately 23°) when the elbow portion of the L-valve remains horizontal. Particles just over the nozzle will continue to flow through until this portion is emptied.

Particles just over the nozzle will continue to flow through until this portion is emptied. Air introduced from the aeration point prefers to fluidize the column of particles in the standpipe rather than push particles over the elbow. Although a higher flow of air will move the powder through the elbow, we avoid interfering with flow settings that may make the powders in the standpipe entrain (bubble out). To restart flow, we preferred to slightly tip the L-valve down (cid:150) a technique that may possibly be incorporated into later designs.159Nozzle diameter(mm)0.5 0.5 0.75 0.75 1.0 1.0Gravity flow or Air assistedG A G A G AAverage mass flow rate (g/s)0.0283 0.045 0.0316 0.0616 0.0717 0.1Volume delivered per second (m3/s)1.9X10-8 3.0X10-8 2.1X10-8 4.1X10-8 4.81X10-8 6.7X10-8Time to draw a 250 (cid:181)m high, 6in.(.1524m) line 0.79 seconds 0.50 seconds 0.71 seconds 0.37 seconds 0.32 seconds 0.22 secondsTable 1.

Results from prototype hopper tests for 100 (cid:181)m glass beads (bulk density 2480 kg/m3).Our mass flow rate results in the third column of Table 1 show that flow rates of a single powder increase with larger orifice diameters. Trials with air entry from the aeration point were conducted at the velocity just below where entrainment in the standpipe was observed. It is believed that the position of the aeration point primarily allows the standpipe(cid:146)s particles to be fluidized, while slightly counteracting the pressure gradient at the nozzle and increasing the flow rate. Flow rate for all diameters were improved by approximately 0.025g/sec with air. Further experiments may reveal whether there is proportional relationship between air-assisted flow and orifice diameter.Array TheoryUsing basic information such as mass flow rate, it is possible to calculate the deposition times of our prototype hopper design its suitability for use with SLS.

To compete with the speed of current roller mechanisms for laying powder on a bed, a successful design should be able to deposit a complete layer in a matter of seconds. The following assumptions characterize the ability of our prototype hopper (and eventually a hopper array) to achieve deposition times appropriate for SLS:Assume that the volume of material deposited per second in m3/s is:V(cid:1)=sG 1( ε ρ − ) B(3)where Gs = mass flow rate of particles (g/s)ρB = ε =bulk density of the powder (g/m3) voidage of powder particles, assumed to be 0.4, based on 60% packing densityAlso assume that a stationary orifice is capable of depositing a disk of material with thesame diameter as the orifice.

Therefore, the volume of the disk (m3) to be filled is:V=π D 42 0h(4)where D0 =hopper orifice diameter (m)h = desired height of a disk of material (m)Therefore the time (s) to deposit a single disk of material with height h is:160V • V== tρπ 2 hD B 0 4 G1(s−ε )A series of these deposited disks laid side-by-side can be thought of as collectively forming a line. With knowledge of the diameter of the disks, the number of disks that compose a line of desired length is the ratio l/D0. The overall time,T, required to deposit a line is:T=tl D0=ρπ lhD 0 B 4 G1(s−ε )The last two columns of Table 1 show that the volumetric flow rate and the time required for drawing a 0.010 inches (250 (cid:181)m) high, 6 inches long line (an estimated minimum length for a powder bed) would be feasible for use with SLS.

Even though delivery times for potentially smaller orifice diameters are above 1 second for gravity flow or air-assisted, an array of nozzles that produces a complete layer in a similar amount of time will be faster than current roller or doctor blade techniques. Shown in Figure 3 are two array concepts to be explored for minimizing delivery time to the bed.Figure 3. (a) A 4x5 staggered array, and (b) A array covering the entire bed length.FUTURE WORKThere are many challenges for future work on a micro-hopper array. It is our aim to determine experimentally and eventually theoretically a prediction for mass flow rate in micro- hoppers. Theory should apply to a variety of powders, so a single highly compatible hopper161(5)(6)design and model for fluidization with dispensing adjustments can be achieved. Visually, the micro-hoppers are delivering powder with consistent flow rate.

Visually, the micro-hoppers are delivering powder with consistent flow rate. However, techniques to aid in understanding all of the phenomena, i.e., determining the pressure at the aeration point and minimum fluidizing velocity in the particle delivery system, need to be determined. Incorporating an XYZ positioning system will yield valuable information on whether our desired minimum feature size for lines and patterns is feasible. Once these goals are achieved, work will focus on the design and control of the array and powder supply system. We look forward to experiments on mixing and depositing powders in gradient fashion with the ability to draw two or more powders from a reservoir supply systemCONCLUSIONSExisting theory on large hoppers was used to guide the design of prototype micro- hoppers for powder delivery. Information on relevant particle properties and flow control techniques was noted.

Information on relevant particle properties and flow control techniques was noted. A prototype micro-hopper design was completed and tested with 100 (cid:181)m glass beads for both gravity flow and gas assisted flow conditions. Mass flow rates were noted to increase with air-assistance, and measured delivery times were found to be appropriate for use with SLS.REFERENCES [1] Beverloo, W.A., H.A. Leniger and J. van de Velde, (cid:147)The Flow of Granular Solids through Orifices(cid:148), Chemical Engineering Science. Vol. 15, pp. 260-69 (1961). [2] Geldart, D., (cid:147)Types of Gas Fluidization(cid:148), Powder Tech. Vol.7, pp. 285-92 (1973). [3] Jenike, A.W., (cid:147)Gravity Flow of Solids(cid:148), Trans. Instn Chem. Engrs. Vol. 40, pp. 264-71 (1962). [4] Jing, S., and H. Li, (cid:147)Hopper Valve for Fine Powders in a Moving-bed Standpipe(cid:148), Powder Technology. Vol. 114, pp. 252-54 (2001). [5] Jing, S., and H.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesign of a Micro-scale Selective Laser Sintering SystemNilabh K. Roy1, Chee S. Foong2 and Michael A. Cullinan11Department of Mechanical Engineering University of Texas at Austin2NXP Semiconductors Austin, TexasAbstract Micro and nanoscale additive manufacturing methods employing metals and ceramics have many promising applications in the aerospace, medical device, and electronics industries. However, the present state of art metal additive manufacturing tools have feature-size resolutions of greater than 100 µm, which is too large to precisely control the geometrical and dimensional aspects of the parts they produce.

The weakness is particularly profound in application of additive manufacturing to the fabrication of fine pitch interconnects in the packaging and assembly of integrated circuits. A new microscale selective laser sintering (µ-SLS) is being developed in this research to improve the minimum feature-size resolution of metal additively manufactured parts by up to two orders of magnitude, while still maintaining the throughput of traditional additive manufacturing processes. This paper presents a detailed design of the sub-assemblies of the µ-SLS system including innovative design features such as the use of (1) ultra-fast lasers, (2) a micro-mirror based optical system, (3) nanoscale particles, (4) a precision spreader mechanism, (5) long-range precision XY nano-positioner and (6) a global positioner.IntroductionIntegrated circuit (IC) package building blocks are very small in dimensions.

Interconnection traces are typically designed to be in the order of 15μm to 25μm in width and 20 μm in thickness. Increasingly finer pitch and feature size to cater for increased IO density, smaller package foot print, and package level integration have fueled industry needs to pursue back end of the line (BEOL) manufacturing processes for package assembly but come with higher costs. From the perspective of feature size and dimensions, there is a gaping disconnect in feature size capabilities achievable using BEOL wafer fabrication processes and organic build up substrate processes. One possible approach to bridge the gap is to develop a low cost intermediary manufacturing process. Laser assisted localized sintering of metallic alloy as used in additive manufacturing stood out as a potential method to fill the role.Additive manufacturing or 3D printing is a process of making a three-dimensional solid objects of virtually any shape from a digital model.

3D printing is achieved using an additive process, where successive layers of material are laid down over previously solidified layers. 3D printing is distinctly different from traditional machining techniques which mostly rely on the removal of material by methods such as cutting or drilling (subtractive processes).A number of additive processes are now available. They differ in the way layers are deposited to create parts and in the materials that can be used. Some methods melt or soften1495material to produce the layers, e.g. selective laser melting (SLM) or direct metal laser sintering (DMLS), selective laser sintering (SLS), fused deposition modeling (FDM), while others cure liquid materials using different sophisticated technologies, e.g.

stereo-lithography (SLA).In our present research, a 3D printing process as practiced in the macro-sized domain has been modified and improved to assemble IC packages offering very high volume through put with potential to revolutionize the entire IC backend assembly industry. The gist of our research and development is to assemble IC packages by fabricating electrical interconnect entities and dielectric build up layer by layer using a micro selective laser sintering process (µ-SLS). The µ- SLS process and machine we are developing works well with many metals and their alloys. We are using µ-SLS to manufacture IC packaging building blocks such as –Inter-layer vias - Conducting traces / wires - C4 bump pads - BGA pads - Flip Chip C4 pillar bumps - AntennasWire bond finger pads - Power and ground planes - Passive ComponentsFigure 1.

IC Package and Assembly Interconnect Building Blocksµ-SLS replaces the semi additive copper plating / etching process typically employed to build up IC package substrate redistribution layers. In addition to building the network of conductive interconnecting layers, the µ-SLS can be used to form electrical connection between the sintered terminals of the conductive traces and interconnection pads on IC chips and between layers. The µ-SLS process described in this paper is a completely additive process where the required interconnect entities in their final form and functions are made at locations where they are needed on the package.The scope of our present research encompasses four distinctly unique but interdependent areas.

In order to accomplish the objectives as laid out above, the team is working on 1) Development of Copper Nanoparticle Inks 2) Physical and Thermal Modeling of Laser-Ink Interaction and Sintering 3) Micro Laser Sintering Process Development and Characterization and 4) Laser and Machine Design and Automation.1496We set out to demonstrate the feasibility of fabricating 3D laser sintered copper pillar bumps on flip chip devices focusing on pillars with high aspect ratio (> 2:1), smaller diameter (< 5µm), and pillar pitch (< 20µm) in near net-shape features. Following that, we extend the process capabilities to enable a low cost and mask-less fabrication process for multi-layer redistribution layer (RDL) with design features comparable to electroplating process for wafer level chip scale device package (WLCSP).

Next, we formulate and demonstrate the new fabrication technology as a feasible low cost and high volume manufacturing alternative for IC interconnect process and for the emerging low cost packaging needs of internet of things (IoT) and wearables markets. And finally, we explore further other potential applications of the fabrication technology in fast IC packaging prototypes, on-chip low R interconnect, customizable IC packages, MEMs sensor design and packaging, and roll-to-roll printed electronics. In this paper, we focus on the design of the laser systems, machine design, and automation aspects of the research.Design of the µ-SLS systemA) Laser and opticsThe first prototype of the µ-SLS uses a 3W CW laser (532nm) from Lasever, Inc. To achieve precise control over the heat-affected zone of the micro-SLS powder bed, the current laser will be replaced with a nanosecond laser, Talon 532-20 from Spectra Physics with a 25ns pulse duration.

Optical components in the setup including mirrors, Digital micro-mirror device (Pro6500 DMD from Wintech), mechanical shutter and objective lens have been designed keeping in mind the possible need of replacing the nanosecond laser with ultra-fast lasers such as picosecond and femtosecond lasers at later stages of product development. Traditionally, for macroscale selective laser sintering system, the optics typically consisted of a set of beam expanders, a few reflection mirrors, and a 3-axis dynamic scanner or an fθ lens and an x-y galvanometric mirror[1]. However, in order to get down to a spot size of 1 μm additional focusing optics are required. One major drawback of focusing the laser spot down to spot of 1 μm from a spot of 100 μm is that the throughput of the system is reduced by a factor of 10,000.

One major drawback of focusing the laser spot down to spot of 1 μm from a spot of 100 μm is that the throughput of the system is reduced by a factor of 10,000. To overcome this problem and increase the throughput of the process, a digital micro-mirror device (DMD), Pro6500 from Wintech, was used to replace the traditional galvanometric mirror in macroscale SLS systems. The micro-mirror pixels are 7 μm by 7 μm in size and are placed 600nm apart[2]. The laser is coupled to the DMD engine via a fiber coupling setup (shown in Figure 1) with a mechanical shutter (VS14 from Vincent Associates) to modulate the exposure time. The incident laser beam into the DMD is then expanded using a fisheye lens and directed off the micro-mirror array through a 10X infinity corrected long working distance objective lens from Mitutoyo. The objective lens has an NA of 0.28 and a theoretical resolving power of 604nm with a Gaussian beam of 532nm wavelength.

The objective lens has an NA of 0.28 and a theoretical resolving power of 604nm with a Gaussian beam of 532nm wavelength. The resolving power of an objective lens for a Gaussian beam is given by equation 1 [3],𝑅𝑒𝑠𝑜𝑙𝑣𝑖𝑛𝑔 𝑝𝑜𝑤𝑒𝑟 =𝜆 𝜋 ∗ 𝑁𝐴where, 𝜆 is the wavelength of the light and NA is the numerical aperture of the objective lens. This allows each 7.6 µm by 7.6 µm pixel in the micro-mirror array to be focused down to a spot size of approximately 1µm. The DMD chipset selected consists of an array of 1920 by 1080 individually addressable micro-mirrors with a binary pattern rate of up to 5 kHz [4]. Therefore, the incident laser beam can be split into a total of 2,073,600 individual beams with a feature size of 1μm each1497( 1 )[5].

Therefore, the incident laser beam can be split into a total of 2,073,600 individual beams with a feature size of 1μm each1497( 1 )[5]. This results in a significant improvement in the throughput of the μ-SLS process over the single beam sinter system.Although the DMD system helps to increase the throughput of the system, it also increases the power requirements for the laser considerably, since a much larger area is being sintered at any given time[1]. In order to size the laser, a simplified model was created to estimate the energy required to sinter one layer of nanoparticles in the powder bed.

In order to size the laser, a simplified model was created to estimate the energy required to sinter one layer of nanoparticles in the powder bed. In this model it is assumed that: (1) the properties of the nanoparticles such as the specific heat capacity and density are the same as that of the bulk material, (2) the particles have been observed to start sintering at 330°C- 450°C[6] which is much lower than their melting point, and (3) all of the laser energy is expended in raising the temperature of the particles—that is, there are no losses in the energy transfer to the surroundings or neighboring particles.

Using this model, the energy required to sinter a spot size of diameter D and layer thickness, h in n pulses, can be found using equation 2 [1], where f is the repetition rate, ρ is the powder density, CP is the specific heat, lf is the heat of fusion, Ti is the initial temperature of the powder bed, Tf is the sintering temperature, R is the reflectivity of the powders and α is the effective power retention factor of the optical elements (mirrors, tube lens and objective) in the setup.𝐸𝑛 =𝜌 ∗ 𝜋 ∗ 𝐷2 ∗ ℎ ∗ 𝐶𝑃 ∗ (𝑇𝑓 − 𝑇𝑖) 4 ∗ 𝑛 ∗ (1 − 𝑅) ∗ 𝛼;Talon 532-20 laser has a rated average power of 20W at a repetition rate of 50 kHz [7]. Assuming that the laser beam loses 90% of its energy till it reaches the powder bed due to losses in the optical elements in the setup, total energy reaching the sample bed would be 2W[1]. Assuming zero time delay in the process i.e.

Assuming zero time delay in the process i.e. sintering can be done successively with zero time delay between two sintering steps, the throughput of the process was estimated to be 594 mm2/s for a layer thickness of 1µm. The theoretical volumetric throughput of the system was thus estimated to be 2.14 cm3/hr. The throughput of this system is two times that of the previous design of the system reported in [1].Figure 2. a) Schematic of the laser-fiber-Wintech Pro6500 coupling b) CAD model of the optical sub-system1498( 2 )B) Use of nanoparticlesμ-SLS requires the use of nanoparticles(NPs) since the particle size needs to be an order of magnitude smaller than the melt pool in order to accurately sinter particles together to form a part. Initial tests performed to spread dry powders resulted in large agglomerates of the nanoparticles forming in the powder bed, as shown in Figure 3a [2].

Initial tests performed to spread dry powders resulted in large agglomerates of the nanoparticles forming in the powder bed, as shown in Figure 3a [2]. These large agglomerates form due to the high van der Waals interactions present between nanoparticles[8] .40nm Average Diameter ParticlesAgglomeratesFigure 3. a) Agglomerated Cu nanopowders b) A spin coated nano-copper ink layer showing uniformity in size distributionDue to these reasons and the fact that spreading dry nanopowders to achieve the packing density for good quality sintered part is very difficult, nanoparticle inks were tested. Solvents in the nanoparticle inks help to reduce van der Waals interactions during the spreading process [6] and allow for a smooth, uniform layer of nanoparticles to be spread onto the build surface, as shown in Figure 3b. Cu NP inks can be easily spread over a substrate with very good packing density using methods like spin coating and slot die coating.

Cu NP inks can be easily spread over a substrate with very good packing density using methods like spin coating and slot die coating. Figure 3 shows samples prepared using Cu IJ70 ink from Applied Nanotech Inc. The ink was spin coated at 1000rpm using EZ6 spin coater from Best Tools, LLC on an Aluminum substrate. The thickness of the layer was measured using a Dektak 6M stylus profilometer and was found to be 430±64nm.C) Spreader mechanism- Slot die coating mechanismIn order to achieve near-net shape µ-SLS features, the layer of Cu NP has to be deposited with uniform thickness with all the particles well dispersed and distributed. Powder bed thickness no larger than 1 um is required for accurate geometry and shape definition when laser sintered. However, it is well known that due to their large surface energy, nano size particles regardless of material types, tend to agglomerate into larger crumps of particles.

However, it is well known that due to their large surface energy, nano size particles regardless of material types, tend to agglomerate into larger crumps of particles. In order to disperse the nano particles into uniform powder bed, surfactants in the form of coatings are regularly used. In the case of our powder bed, we have found that amorphous nano copper powders difficult to be spread over a large surface area with well controlled thickness. As a result, we have resorted to use solvent based copper nano powder inks to form our powder beds for sintering.There are many different processes available to lay down a thin layer of ink.

As a result, we have resorted to use solvent based copper nano powder inks to form our powder beds for sintering.There are many different processes available to lay down a thin layer of ink. Among them are spin coating, doctor blading, screen printing, ink jet printing, pad printing, slot die coating, gravure printing, curtain coating, spray coating, flexographic printing, and offset printing.1499Slot die coating method has been adopted because it is a very flexible and adaptable process highly suitable for the development work we are undertaking. Such a slot die coating system is advantages because it is able to coat organic or inorganic liquids on flexible or rigid substrates with excellent coating uniformity (typically ±3% or better). The slot die coating system we are working on is able to deposit a wide range of thicknesses ranging from 20nm to 150μm.

The slot die coating system we are working on is able to deposit a wide range of thicknesses ranging from 20nm to 150μm. Our slot die coating system is also capable of coating a wide range of process materials with varying viscosities.From the high volume manufacturing perspective, the slot die coating process is also very attractive because it has high efficiency with over 95% material utilization.

In addition to being a very high throughput, the slot die coating process can be easily adapted for roll-to-roll processing which will further improve the overall productivity.Slot die coating technology has been used to deposit a variety of liquid chemistries onto glass, stainless steel, and plastic substrates for the development and production of a broad range of applications such as flat panel displays (including LCD, OLED, flexible, and others), photovoltaic panels (such as CIGS, CdTe, OPV, and others), solid state lighting (OLED and LED), and various organic and printed electronic applications such as RFID, polymer batteries, sensors and others. This is achieved through the precise metering of the process fluid and dispensing it at a controlled rate while the coating die is precisely moved relative to the substrate.Figure 4.

This is achieved through the precise metering of the process fluid and dispensing it at a controlled rate while the coating die is precisely moved relative to the substrate.Figure 4. Slot Die Coating Working Principles[9]Slot-die coating is a non-contact large-area processing method for the deposition of homogeneous wet films with high cross-directional uniformity. It can handle a broad range of viscosities between 1 mPa⋅s and several thousand Pa⋅s while the coating speed has a similar wide spectrum between 1 m/min and 600 m/min. The working principle is shown in the figure above. The wet film thickness is controlled by the flow rate, coating width, and speed.

The working principle is shown in the figure above. The wet film thickness is controlled by the flow rate, coating width, and speed. The resulting dry layer thickness d in cm for a given ink can be expressed with the formula𝑑 = 𝑓 ∗𝑐 𝑣 ∗ 𝑤 ∗ 𝜌;where f is the flow rate in cm3/min, v is the coating speed in cm/min, w the coating width in cm, c is the concentration of the solids in the ink in g/cm3, and ρ the density of the material in the final film in g/cm3[10].A slot die coating system comprises of three main components:1500( 3 )1. Slot Die AssemblyThe liquid ink is pre metered by the slot die onto the substrate subjected to the fluid rheology, pump design, and die lips geometrical construct. In a typical construction, the slot die is made of two separable halves in which a liquid reservoir is machined into. Both halves are bolted together with a metallic shim between them to provide the slot gap.

Both halves are bolted together with a metallic shim between them to provide the slot gap. Enclosed within these two halves are flow channels machined with ports for liquid inlet and recycling. The major feature of the die set is the reservoir that opens to its widest segment corresponding to the coating lips width. The reservoir distributes the coating fluid that enters the die to its full target width and is designed to generate a uniform, streamlined flow of material through the exit slot of the die.Figure 5. A slot die set showing the reservoir and lips where ink to be coated exit[11]2. Slot Die Positioner and Substrate CarriageAn adjustable Z height track is used to hold the slot die assembly and to precisely position the slot die at the optimum angle and proximity above the substrate surface. The substrate held down by the chuck is moved by the carriage below the positioned slot die as inks is flowing out of the slot of the die.

The substrate held down by the chuck is moved by the carriage below the positioned slot die as inks is flowing out of the slot of the die. The slot die positioner is important for stabilizing the interaction between die and moving web utilizing the angle of attack between die and substrate, the distance between the two and the degree of offset between the lips. These parameters can be adjusted to optimize the process. The entire assembly must be isolated from vibrations to ensure precise and uniform coating deposition.3. Fluid Delivery SystemThe fluid delivery system has to provide a non-pulsing, continuous feed of ink to the die in a controlled pressure. In addition, excess ink is flowed back to the ink reservoir and will be feed back into the slot die.

In addition, excess ink is flowed back to the ink reservoir and will be feed back into the slot die. This critical piece of equipment works in combination with an accurate line speed control to determine coat weight of the fluid.The design and fabrication of the slot die set are critical to the optimal and precision operation of the coating system. The smoothness and flatness of surfaces at or near the exit slot of the die are crucial for product quality. These surfaces include the lip land just inside the exit slot and the face of the wet lip (the lip passed by the moving web after it has passed the exit slot). Improvements in dimensional tolerances made possible by precision machining of these surfaces1501may seem miniscule, but in applications involving high volumes or costly coating materials, the economic gains can be great.

The effect of flatness variations becomes progressively greater as coating weights become smaller.The slot die coating method offers higher line speed capability and greater coat weight uniformity provided that the coating system as a whole must be designed to exacting tolerances based of the rheological characteristics of the fluid. The slot die can then achieve defect-free coatings with increased production speeds, positive coat weight control, controlled cross-web distribution and improved waste management.D) Long range XY nanopositionerA long range XY nano-positioner is employed to align the sample wafer under the objective lens of the optical system and to scan the sample under the micro-mirror based optical system with sub-micron accuracy and repeatability. Due to absence of friction and backlash, flexure-based compliant guiding mechanisms have been widely employed in applications requiring precision motion.

Due to absence of friction and backlash, flexure-based compliant guiding mechanisms have been widely employed in applications requiring precision motion. The micro-mirror based optical sub-system can only pattern an area of 2mm by 1mm at once and thus, the wafer has to be moved under the sub-system to pattern the entire 2” by 2” area. Consequently, the need to have the range of the system to be 2” along X and Y directions. Selection of an actuator for this application is vital to achieve the 2” range with sub-micron accuracy. Electromagnetic actuators or voice coil actuators (VCAs) are widely used in long-travel, flexure- based parallel XY stages since VCAs are capable of delivering a frictionless inch-level stroke along with a fine resolution.Figure 6.

A double parallelogram flexure mechanism (DPFM) unit[12]Although the 2” stroke can be generated by selecting suitable actuators, it is challenging to devise a flexure parallel stage with a large range and compact size simultaneously. To accomplish a large motion range, slender leaf flexures are usually employed. The minimum width of the flexure is restricted by the practical tolerance of manufacturing and the maximum length is constrained by the compactness requirement [13]. Double parallelogram flexure mechanisms (DPFMs) has been used in the design to achieve the long range motion. Assuming the axial force and tangential displacements of the flexure beams to be zero, the maximum displacement of the flexure beam is given by equation (3) where xmax is the maximum unidirectional displacement of the DPFM unit, h is the thickness of the flexure beam and FOS is the factor of safety.𝑥𝑚𝑎𝑥 = 𝜎𝑦 ∗𝐿2 3 ∗ 𝐸 ∗ ℎ ∗ 𝐹𝑂𝑆1502( 4 )Figure 7a) shows the design for the flexure stage.

The output motion is guided by four DPFMs located at the corners of the module. The second module is mounted orthogonally on the top of the first module, to get a parallel-kinematic and stacked structure. This modular design allows the cost reduction in both manufacturing and maintenance. When driven by VCA1, four DPFMs on the top and two DPFMs at the bottom guide the motion along X and similarly for the motion along y direction. Hence, a decoupled two-dimensional translation is achieved by the XY stage.Figure 7. a) XY nanopositioning sub-assembly b) Static FEA simulation results- x-displacement with an input Fx = 1NThe stiffness of the stage can be modeled as a combination of six springs in parallel for each direction and hence, the stiffness value can be estimated by𝐾𝑥 = 𝐾𝑦 = 6 ∗12 ∗ 𝐸 ∗ 𝐼 𝐿3; =6 ∗ 𝐸 ∗ 𝑏 ∗ ℎ3 𝐿3Using the estimated values of stiffness and range, the actuation force needed to obtain the desired range was estimated to be 16.53 N.

The FEA simulation conducted (see Figure 7b) using SolidWorks gives a stiffness value of 838.9 N/m which is higher than the stiffness obtained using equation (5). This could be due to two reasons- 1) adding fillets for ease of water-jetting the flexure increases the stiffness and 2) approximation of zero axial force and tangential displacement doesn’t hold true. Using the stiffness value obtained from FEA simulation, the actuation force for obtaining 1” travel was calculated to be 21.31N. Based on the range and actuation force value, a suitable VCA- NCC34-25-060-1X from H2W Technologies with a maximum travel of 3.4” and a continuous actuation force of 38.7N was selected for the stage. Figure 7a shows the CAD model of the long range XY nanopositioner. A polymer plate with a viscous fluid layer on both sides is sandwiched between the two flexure plates to damp out the vibration of individual DPFMs.

A polymer plate with a viscous fluid layer on both sides is sandwiched between the two flexure plates to damp out the vibration of individual DPFMs. The modal analysis (see Figure 8) of the structure shows the first two modes to be associated with X- axis and Y-axis motion of the center-stage. The center stage has actuators (VCAs) attached to it and hence, these modes can be controlled using an appropriate control action. The higher modes of the structure are associated with the vibration of individual DPFM units which have no actuators attached to them and thus, they need to be damped passively. The viscous fluid between the polymer plate and the flexure stage aids in passive damping of these higher modes.1503( 5 )Figure 8. Modes of the XY stage- obtained using Frequency analysis in SolidWorksThe current setup utilizes a step-and-flash system where an image is formed on the DMD and that image is transferred to the particle bed.

The entire powder bed is then moved over by the width of the image using the XY stage and the process is repeated.Table 1. Structural parameters, stiffness and range of the stage using properties of Al-7075ParameterValueParameterValuel101.4 mmKx651 N/mh0.5 mmKy651 N/mb12.7 mm x_max24.14 mmFOS2Total range48.28 mmE) Global travel between slot die coating and optical stationsFor the global travel between the sintering station and the slot die coating station, air bearings linear guide rail have been employed in the system driven by a direct-drive linear servo motor. Since the desired feature size is 1µm, it is required that the registration errors under the optical station and slot die coating station are under ±0.5 micron and the motion between the two stations is highly repeatable and free of stick slip phenomenon. Other actuation mechanisms such as a lead screw with stepper motor or a ball screw with stepper motor use rolling bearings.

Other actuation mechanisms such as a lead screw with stepper motor or a ball screw with stepper motor use rolling bearings. Although rolling bearings have much lower coefficient of static friction than plain bearings, their static coefficient of friction is higher than their coefficient of kinetic friction which is the major cause of the stick slip phenomenon experienced with actuation mechanisms employing plain and rolling bearings. Air bearings have many more advantages over rolling bearings including near zero friction and wear, high speed and high precision capabilities, and no oil lubrication requirements. Friction in air bearings is a function of air shear from motion, so at zero velocity there would be zero friction making infinite motion resolution theoretically possible[14] with minimal hysteresis or reversal error.

As a result, air bearings are highly repeatable and ideal for many inspection and manufacturing operations, such as flat-panel inspection, semiconductor metrology, and laser machining[15]. A direct-drive motor and high-resolution encoder can position a moving carriage1504supported by an air bearing to within 1 nm in a linear application depending upon the resolution of the encoder. In addition the reduction in the number of components used reduces servicing work and lowers maintenance costs. Air bearings virtually eliminate stiction, which improves resolution capabilities. A direct drive linear servo motor has been used to drive the air bearing trucks on the guides. For displacement sensing, an interferometric sensing system from Attocube with a resolution in the range of fraction of a nanometer has been employed to achieve very fine resolution on the overall travel.Figure 9.

Air bearing sub-assembly with XY nanopositioning stageF) Chuck for holding the waferCu nanoparticle inks have been observed to have an initial exothermic peak near 180- 190°C which is a result of the removal of residual solvent left in the sample after the sample has been spin coated and dried[6]. Adding a heater to the stage to heat up the sample to an elevated temperature to remove the solvent is one way to get a uniform sintering across the sample. This heating element has the additional benefit of reducing the temperature gradient between the sintered and the un-sintered areas, which helps to reduce warpage and thermal stresses in the part.Although adding the heating element to the chuck aids in achieving a uniformly sintered part, the heater may also lead to development of thermal stresses in the chuck causing warpage in the top chuck.

This warpage of top chuck affects the quality of slot die coating as the gap height between the slot die head and substrate changes due to thermal expansion of the top chuck. The material and geometry for the top chuck thus becomes critical to eliminate any excessive thermal warpage of the top. C145, a Tellurium alloy of Copper(Cu) was selected as the material for fabricating the chuck over Aluminum(Al) and Copper due to its lower coefficient of thermal exapansion-17.1 ppm/°C against 24 ppm/°C and 23 ppm/°C for Al and Cu respectively. These predictions were verified by FEA simulations conducted using SolidWorks. Kapton heaters were added to the back side of the top plate of the chuck which can generate a heat flux of 30W/in2. The ball plungers press against the top plate to hold it, however it also allows the plate to expand in plane so that the wafer holding region doesn’t warp under thermal stresses.

The ball plungers press against the top plate to hold it, however it also allows the plate to expand in plane so that the wafer holding region doesn’t warp under thermal stresses. However, if the top plate were bolted to the bottom plate, the out of plane displacement of the top plate will be significantly higher (refer to Figure 11a and 11b). To get the out of plane displacement of the top plate of the chuck, all the plates were heated to temperatures 356K using the Kapton heaters’ thermal flux. These temperature profiles were then used to get the out of plane displacements of the top plates (shown in Figure 11). The relative out of plane displacement of the design where the1505top plate is bolted to the bottom with C145 as the material, was estimated to be 1.4µm against the 0.9µm relative displacement obtained with the design with ball plungers in place using the same material (refer to Table 2).

Table 2 lists the maximum relative displacement found out across the diagonal line shown in Figure 11a). This is the relative out of plane displacement of the top chuck plate and should be minimized to get uniform coatings of Cu nano-ink.Figure 10. Chuck design for holding the sample waferFigure 11. a) Material- C145, with the top plate bolted to the bottom plate b) Material- C145, c) Material- Al and d) Material-Cu with top plate constrained using ball plungers as depicted in Figure 101506Table 2. Table showing relative out of plane displacements obtained with designs shown in Figure 11a)1.4b)0.9c)2.1d)1.3Designs in the order as shown in Figure 11Relative out of plane displacement (μm)ConclusionCAD model of the assembly has been shown in Figure 12.

The design of the microscale selective laser sintering system presented in this paper will give us the ability to start to produce true three dimensional parts in metals with microscale feature sizes and thus, control the microstructure of materials. Through the use of micro-mirror array based optics and high speed, precision nanopositioning systems it is possible to produce parts with micron scale features at a theoretical build rate of up to 2 cm3/hr. This system provides a revolutionary approach to IC packaging as it eliminates the use of the conventional lithography processes in such as masking, exposure, plating and etching and thus reducing the cost and manufacturing time for ICs.

Its ability to produce additively-manufactured parts with controlled microstructures (which allows us to tune the properties of materials such as density, strength to weight ratio, elastic modulus, fracture toughness and so on) has numerous applications in medical, automotive, aerospace and many other secondary industries.Figure 12. CAD model of the current µ-SLS system1507References[1] Roy, N. K., and Cullinan, M. A., “μ -SLS of Metals: Design of the powder spreader,powder bed actuators and optics for the system,” pp. 134–155.[2] Roy, N., Yuksel, A., and Cullinan, M., 2016, “DESIGN AND MODELING OF AMICROSCALE SELECTIVE LASER SINTERING SYSTEM,” Proceedings of the 11th Manufacturing Science and Engineering Conference MSEC16, Blacksburg, Virginia.[3] Hecht, E., and Zajac, A., 1974, Optics Addison-Wesley. [4]Imaging, M., Measurement, S., and Lighting, A., 2015, “DLP5500 DLP ® 0 . 55 XGA Series 450 DMD,” (1).

isign of a SoliidFreefotm· Fabrication Diamond Reactorw. Richards Thissell,1,2 James Tompkins,2 and Harris L. Marcus1,21. Department of Mechanical Engineering 2. Center for Materials Science and Engineering ETC 9.104 University of Texas at Austin Austin, Texas 78712 (512) 471-1504 FAX: (512) 471-7681 AbstractSolid Freeform Fabrication (SFF) has progressed from the visualization aided stage of computer aided designs (CAD) to rapid prototyping of structural parts. Among the promising techniques for producing structural prototypes is the tech nology of chemical vapor deposition (CVD) of polycrystalline diamond. This paper discusses the thermodynamic and kinetic theories that suggest that structural diamond may be rapidly deposited at rates approaching 1 mmJhr from the vapor phase at metastable thermodynamic conditions. The design of a reactor that will produce structural diamond prototypes is discussed.

The design of a reactor that will produce structural diamond prototypes is discussed. This reactor combines down stream microwave plasma enhanced chemical vapor deposition (DMWPECVD) with a scanned CO2 laser that locally heats the substrate to diamond deposition temperatures. The input:Fases are H2, 02' CH4, and Ar. The operating pressure range to 7 x 102 Torr. The reactor is designed for in situ deter of the reactor is 1 x 10- mination of deposit thickness while deposition occurs as well as having the capacity of fitting on an existing resonance enhanced multiphoton ionization time of flight mass spectroscopy (REMPITOFMS) apparatus that will allow for plasma diag nostics immediately above the heated substrate.

~ lee Net Rate of Change of species i, sec·1 Sticking Coefficient, Dimensionless :BoltzmannConstant,K= 1.380x 10-23i Wavelength, m Viscosity, Pasec Density of speciesi, k~ A Lennard-Jones Parameter, A Time Required To Approach Steady State, sec Fuller, Schettler, and Gidings Diffusion Volume Incrementsm2 History of Vapor Phase Diamond SynthesisDeposition of solid material from selected gases has great .potential as an alternative freeformJabrication technique. The use of reactant gases (or precursors) in a chemical reactor to grow layers of a desired material is commonly known as Chemical Vapor Deposition (CVD), and is widely on a commercial scale. An exciting and unique application of this concept is used below to propose the freeform fabrication of diamond from gas phase hydrocarbons.Diamond growth from the gas phase for freeform fabrication is complicated by factors not encountered in normal chemical vapor deposition.

First, diamond is a metastable phase at all reasonable temperatures and pressures. Under most conditions, carbon will be. deposited having the equilibrium sp2 bonding. However, under kinetically controlled conditions, l some carbon is deposited in the metastable sp3 bonded state. The activation energy needed for dia- mond to transform to the more stable graphite structure is quite high at 730 ± 50 ~l for {110} planes ,2 making diamond quite stable at temperatures below 1600 °C in oxygen partial pressures less than 10-6 Torr. removal of sp2 bonded carbon may be accomplished during growth by using a high percentage of atomic hydro;enin the gas phase, which preferentially etches sp2 bonded carbon over sp3 bonded carbon.

removal of sp2 bonded carbon may be accomplished during growth by using a high percentage of atomic hydro;enin the gas phase, which preferentially etches sp2 bonded carbon over sp3 bonded carbon. Furthermore, hydrogen stabilizes the dangling sp3 bonds present at the surface of the diamond layer and prevents them from collapsing into a sp2 bonded surface structure.The first reported growth of diamond from the gas phase was from von Bolton in Germany (1911), using acetylene and mercury at 100 0 C.4 Simple acetylene torches were used in Germany by researchers beginning in the 1940' s. 5 The first documented work confirmed by others was by Eversole of Union Carbide in 1952.6 It is interesting to note that this is two years earlier than the first diamond synthesis by the well known high pressure method developed by General Electric in 1954.7 In 1956, a patent was filed in the Soviet Union by Deryagin and Fedoseev for the growth ofdiamond from carbon containing gases.

8 Both Eversole and Deryagin deposited graphite and diamond simultaneously on the substrate, then preferentially etched away the graphite with oxygen or high pressure hydrogen to give a small net gain of diamond. In 1966, Lander and Morrison of Bell Telephone Laboratories discovered the importance of a hydrogen-rich environment in preventing the conversion of the diamond surface to graphite due to unsatisfied sp3 bonds. 9 Later, Angus from Case Western Reserve University used this information and Eversole's hydrogen etching to alternately grow and clean diamond films. 1O He also increased the growth rate by using a methane and hydrogen gas mixture. Finally, Vickery of Diamond Squared Industries combined the growth and etching cycles into a con tinuous, single step process by using 95% hydrogen and 5% hydrocarbon (1978).1150The key to rapid, continuous film growth was provided by Deryagin.

His identification of atomic hydrogen as a rapid graphite scavenger dictated the design of all subsequent diamond growth chambers. He describes l2 five methods of generating astable source ofatomic hydrogen - catalysis, ion beam, electron beam, heated filament, and plasma excitation, the latter two of which are used on nearly all modem reactors. Deryagin also. demonstrated continuous diamond film growth on a non-diamond substrate, a milestone in the pursuit of a commercially viable process.1974, scientists from Japan's National Institute for Research in Inorganic Materials (NIRIM) began work on diamond deposition from the gas phase. Using the heated filament method above, they produced films more suitable for commercial use. Later, glasma based methods made their coatings easier to produce and more consistent in quality.

Using the heated filament method above, they produced films more suitable for commercial use. Later, glasma based methods made their coatings easier to produce and more consistent in quality. Japan now holds by far largest number of patents (over 500) on commercial diamond films, and there are many coordinated industrial groups actively pursuing research.3 Diamond Characterization and ClassificationCarbon may bond by several different hybridized orbitals, sp1, sp2, and sp3. Carbon may be amorphous, having either sp bonded carbon (carbyne), sp2 bonded carbon, or sp3 bonded carbon. Carbon may have a hexagonal lattice with.space group P~/nvnc and P63mc for sp2 bonded lattice with space.group P63/mmc (londaleite), and a bonded carbon, a hexagonal bonded crystal lattice with space group Fd3m. Carbon solid3 may also be formed diamond with of two or more of the above lattice types. Diffraction data for all of the above groups may be very similar and easily misinterpreted.

Carbon solid3 may also be formed diamond with of two or more of the above lattice types. Diffraction data for all of the above groups may be very similar and easily misinterpreted. Raman spectroscopy lattices and the best technique for characterization of carbon solids. has proven to Raman spectra provide information about the crystal structure, only the electronic state of the not pure graphite consists of only sp2 bonding and that pure diamond consists bonding.only sp3 bonding and that amorphous carbons may. consist of both sp2 and sp3bonding.spectrum ofhexagonal!iraphite consists of a single narrow peak at 1580 em-I, and weak lines at 2450cm-1 and 3250 em-I. 4 The strong 1580 em-I peak i~ due to the sp2 bonding. The diamond crystal lattice exhibits a sharp .peak near 1332 cm-I.IS The Raman scatterin& efficiency for the sp2 1580 cm-l peak is 55.6 times greater than that for the sp3 1332 cm-l peak.

The diamond crystal lattice exhibits a sharp .peak near 1332 cm-I.IS The Raman scatterin& efficiency for the sp2 1580 cm-l peak is 55.6 times greater than that for the sp3 1332 cm-l peak. A substance with a mixture of sp2 and sp3 bonding will have both the 1580 em-I and the 1332 cm-l The peaks will also broaden as the degree of disorder increases. Diamond materials by the ratio of the 1332 cm-l peak to the 1580 cm-l peak and by the are currently full width half maximum (FWHM) of the 1332 cm-l peak. Diamond materials may also have constituents that also affect their properties. There are six diamond non-carbon classifications for natural diamond. Each of these divisions is based on different physical properties that are a result ofimpurities and second phases. Table 1describes the natural diamond types.TheHigh temperature, high pressure synthetic diamonds are similar to Type 1b naturaldiamonds.There are two main types of hard carbon films: diamond-like carbon (DLC) and diamond.

Diamond-like carbon films are generally amorphous, but they can also have crystals about 2-20 DLCcontains from -0-50 mole % hydrogen. DLC films behave like diamond nm in Their filnl hardness can when the carbon gram-atom number densities are above 0.2em3approach that of diamond films and their amorphous character results in smooth coatings.51Type1a Ib 2a 2b Carbonados BallasTable 1 NaturalDiame>nd Types Optical TransnlissignTheI111al Threshold conductivity ;K -9 -9 -26 -26 Variable VariableNatural PrevalenceResistivity nA(nm)> 1.0 X 1016 > 1.0 X 1016 > 1.0 X 1016 10-100098 % 1 % < 1 % 1%320 320 225% polycrystalline 0.2 %polycrystallineImpurity0.1%N <0.05 %N «0.05%N B Variable VariableSecond PhaseNone None None None Graphite and others NoneDiamond films are defined as single-phase films that are single crystals or with grain sizes greater than 50 nm, if polycrystalline.

The film should have a crystalline morphology demonstrating either the four fold symmetry of {IOOf or {110} surface planes or five-fold symmetry of multiply twinned {Ill} lattice planes. 7 SEM analysis should reveal cubic, cubo-octahedral, octahedral, pentahedral (evidence of. multiply twinned {Ill} planes), or rounded cauliflower shaped surfaces for .polycrystalline films. X-ray or electron diffraction would reveal a single-phase crystalline structure. The Raman spectra must consist of a single, distinct, narrow line at about 1332 em-I. Low temperature, low pressure synthetic polycrystalline diamond films are similar to carbonados and ballas in their properties. Boron doped (1020 ato;)emp-type films have been made with resistivities of 10-3 be made with resistivities ranging over 19 orders of magnitude. Arsenic doped films show n-type semiconductivity. P-njunctions are therefor possible with diamond.

Arsenic doped films show n-type semiconductivity. P-njunctions are therefor possible with diamond. High quality diamond films have excellent optical properties with a high index of refraction and are optically transparent from 225-2000 nm (most excimer, He:Ne, and Ng: YAG lasers) and above 6000 nm (C02 lasers).{1. This means that diamond films may emAn exciting new develo~ment is the productIon of high purity 12C diamond films containing about 99.9% 12C.18,1 ,20 Normal diamonds have about 1% l3C. The increase in mole fraction of 12C causes a 50% increase in thermal conductivity and about an order of magnitude increase in laser damage tolerance in diamond.

The increase in mole fraction of 12C causes a 50% increase in thermal conductivity and about an order of magnitude increase in laser damage tolerance in diamond. Calculations indicate that a 100% 12C defect free single crystal diamond may have a thermal conductivity of about 48 W Unpublished research indicates that polycrystalline CVD 12C diamond films do not have a higher thermal conductivity than natural isotope abundance CVD diamond films.524 Mechanism of Vapor Phase Diamond SynthesisMetastable diamond films may be prepared at much lower temperatures and pressures than equilibrium conditions would require. This is a kinetically controlled phenomena.

This is a kinetically controlled phenomena. The exact chemical mechanism of vapor phase diamond deposition is currently unknown.Many different vapor phase processes can lead to diamond deposition including ion beam deposition, electron beam sputtering of graphite, heated filament techniques, slightly oversa turated (carbon rich) oxyacetylene torches, atmospheric plasma torches, various plasma glow discharge deposition processes including DC,21 RF,22 and MW23 plasmas. Laser enhanced chemical vapor de~sition (LECVO) has been attempted by Kitahama et aI., but so far it has been unsuccessful. 4,25,26 Many of these techniques use. different reactant compounds yet result in microstructurally similar diamond deposits. All of these methods involve vapor phase con ditions that result in many reactions becoming significant, leading to very complicated vapor phase compositions. Substrate temperatures for most of these methods must be 800-1000 °C for sp3 bonded carbon to be the dominant phase.

Substrate temperatures for most of these methods must be 800-1000 °C for sp3 bonded carbon to be the dominant phase. Atomic hydrogen is present in all of these methods. Research on determinin~vawr phasediamond precursors is very active and the literature results are often contradictory.,28, 9The diamond deposition reaction may be treated in an overall fashion and significant insights may still be made. 30 Consider a generalized deposition reaction of some unknown gas phase hydrocarbon radical that can react to fonn diamond, graphitizable carbon (true well ordered graphite formation thermal temperatures in excess of 2000 0 C) and atomic hydrogen:k",3Cv~ ~ vCsp3 +wH.k',p3k"'2CxHy ~ xCspz + yH·k $p2The net rates of deposition, assuming an arbitrary reaction order, are given by: aCsp3nanb k sp3P H.,Rsp3 = T = k sp3P C."Hwacspznd Rspz =T =kspzPCxHy - k spzPH· where the rate constants have the customary Arrenhius temperature dependencies.

Gaseous carbon does exist in the thermal equilibrium plasmas and heatedfilament techniques, and may exist in high electron density non-isothermal plasmas. Consider also the general equilibrium reaction between these unknown hydrocarbon radicals and gaseous carbon immediately above the substrate:nc,Cv~ ~v· C ·+wH·CH ~x·C·+yH·xy53[Ia][Ib][2a][2b][3a][3b]having an equilibrium constant:CyHWK. = ·c·P~ PH. ·c·PCyHWCxHyK. = ·c·P~ PlI. ·c·PCxHyWe can also write chemical equations describing the equilibrium gaseous carbon partialpressures over sp2 and sp3 bonded carbon:C 3 ¢::::> ·C .spThe equilibrium gaseous carbon partial pressure over sp2 bonded carbon is twice that over sp3 bonded carbon, according to Badzian et a/.: 31cfp3P~ ·c·= 2 Csp2P~ ·c·Consider a standard gas solution over a polycrystalline solid surface containing both sp3 and sp2 bonded carbonregions. Equation [4] is not satisfied because ofequation [6].

Equation [4] is not satisfied because ofequation [6]. Draw an analogy of the solid surface to an ideal solution, for purposes of interaction with the vapor phase. Then the degree of supersaturation of gaseous carbon is twice as high over the sp2 bonded carbon regions than it is over the sp3 bonded carbon regions: P. ·c· . Csp3P~ ·c·P. ·c·= 2Csp2P~ ·c·We can consider a pseudo equilibrium state by adding reactions [1] and [5] together, which results in:CJ~ ¢::::> v· C . +wH·C H ¢::::> x . C . +yH ." We can write a pseudo equilibrium constant for this equation having the form:yCyH,.K~ltudo =Csp3PO vpw H· . -c- _c-PCyH,.where P~ H and P~ H are determined by setting the rate equation [2] equal to zero and solving:Y W.x..,. 54[4a][4b][5a][5b][6]m[8a][8b][9a][9b]sp3pOCvH.,=[k' k sp3sp31 ];;III" pili.

54[4a][4b][5a][5b][6]m[8a][8b][9a][9b]sp3pOCvH.,=[k' k sp3sp31 ];;III" pili. H·sp2pO = C;cH,[k' k sp2sp21 ];;lidpile: H·is constant for a given set of operating conditions, making the degree of gaseous carbon supersaturation over sp3 bonded and sp2 bonded carbon a function of Pc u and P~ H • If wev''''w% yassume that iK. =iKP:eudo, then the degree of supersaturation becomes:c·c·...........JL..,'" ,,+JL analysis of the power coefficients ofPHe will leadto the following conclusions. If nb/vna > nixnc , then as P He increases, the nucleation and growth ofsp2 bonded carbon would be favored over that of sp3 bonded carbon, which is clearly not the case. Likewise, ifnb/vna = nixnc, then for all values ofPHe the nucleation and growth of one type of carbonis favored over the other.

Likewise, ifnb/vna = nixnc, then for all values ofPHe the nucleation and growth of one type of carbonis favored over the other. the only case that compares favorably to experimental evidence is nb/Vf'la nixnc• Considering this latter case, it can be seen from the equation [11] how sp3 bonded <carbon is favored at higher values of PHe and sp2 bonded carbon· is favored at lower valuesofPHe for constant P . . Therefore, high rate preferential sp3bonded carbon deposition requires highc·gaseous carbon supersaturation over the surface and a high atomic hydrogen partial pressure to inhibit net sp2 bonded carbon nucleation and growth.has another role as a stabilizer for the sp3 bonded carbon surface, which contains an unsatisfied sp3 bond. 32 H· will readily bond to this unsatisfied surface sp3 bond and prevent the surface layer from collapsing to the more stable sp2 bonding in the temperature range of 900-1000 °C.

32 H· will readily bond to this unsatisfied surface sp3 bond and prevent the surface layer from collapsing to the more stable sp2 bonding in the temperature range of 900-1000 °C. 33 Vapor phase halogens will also stabilize the surface unsatisfied sp3 bond. Therefore, keeping the sp3bonded carbon surface from restructuring is an important step tpat inhibits sp2 bonded carbon nucleation and growth and allows for further· epitaxial growth of sp3 bonded carbon.The presence of oxygen in the vapor phase has been shown experimentally to greatly increase the preferential deposition of sp3 bonded carbon. The above analysis does not consider the role 9f oxygenQn the.degree. of sllpersaturatiotlofg~seousCarbonoversp3bo~~edandsp2 bonded· carbon.•.However,itis wellknownthat·sp3 bonded carbon is much·. moreo,,-idation resistant than sp2 bonded carbon. 34 Perhaps the role of oxygen is tp increase the etch rate of sp2 bonded carbon.

-k'spJP~~ -k'o..l:''113Perceiving that k'o > k'o ,one can see from equation [13] that the presence of oxygen in the diamond depositing atmosphere will allow the partial pressure ofthe hydrocarbon to be increased for the same deposit quality.'II''1135 Transport Phenomena In LCVD and LECVD SystemsThe phenomena of momentum, heat, and mass transport are critically important in the proper design of a real LCVD or LECVD system and in understanding the resulting deposit phase, microstructure, and material properties. High rate deposition may be a transport limited process. The gas phase reactants and reaction products must flow through the reactor. The laser beam heats a small region ofthe substrate or previous deposit at any instant in time. This heat is dissipated through heat conduction, convection, and radiation. The reactants and reaction products must be transported to and from the.reaction zone through the combined effects of diffusion and momentum transport.

The reactants and reaction products must be transported to and from the.reaction zone through the combined effects of diffusion and momentum transport. In general, the resulting partial differential equations for diffusion, momentum,continuity, and heat conduction are coupled together and therefore must be solved simultaneously. The standard assumption of negligible convective and radiative heat transfer may be then tested by comparing the resulting conduction heat flux with the heat flux from convective heat transfer correlations and black body radiation.5..1 Photon HeatingProper control of laser induced substrate heating is critical for solid freeform fabrication of diamond. The thermal processing window for diamond deposition is between 800-1100 °C. Light energy can result in thermal heating of a media through absorption.

The thermal processing window for diamond deposition is between 800-1100 °C. Light energy can result in thermal heating of a media through absorption. The general equation for heat flow by conduction into semi-infinite slabs neglecting heat losses due to conduction, convection, and radiation to the surrounding medium is given by:aTPi(T)Cp(T)a;- V[K(T)VT] = Q(t)Q(t) = I~(T,A,e) [15] where Q(T) is the heat source from a laser beam incident on the surface. Consider the case of a laser beam incident on a gas-solid interface oriented perpendicular to the beam. At time zero the beam is turned on.

Consider the case of a laser beam incident on a gas-solid interface oriented perpendicular to the beam. At time zero the beam is turned on. Herziger35 solved the transient heat conduction equation for the center56[l2a][l2b][l3a][14]ofthe beam assuming a Gaussian beam intensity distribution, temperature independent material properties, and no phase transition latent heat effects:~aJ - ~T(O, 0, 0, t) = {f arctan AI.,rf 2K ~ 2 The assumptions of constant material properties and zero latent heat made to develop the closed form solution severely limit its applicabili~ to real-world situations because CpandK can change over an order of magnitude over a 1000 C temperature range. Accurate mOdeling of photon heating requires open form, iterative numeric solutions.36Substrate and Deposit Vapor Pressurelocal substrate or deposit vapor pressure is only a function of the local temperature of the material.

Relations between vapor pressure versus temperature have an empirical nature. Such vapor pressure relations often have of form:lnpsat = A - BC +DT +E In TT+where A, B, C, D, and E are constants. When D = E = 0, equation [15] is called the Antoine equation. When the vapor pressure of the material within the heated zone equals or exceeds the system pressure, rapid local va~orization of material proceeds. Figure 1 shows experimental gaseous carbon vapor pressure3 and a least squares fit to equation [15] assuming that C=500.substrate or deposit vapor itself has an absorption coefficient that can be non-trivial. Therefore, vaporization of material can and usually does increase the coupling of laser energy into the substrate or deposit by absorbing both incident and reflected laser energy.

Therefore, vaporization of material can and usually does increase the coupling of laser energy into the substrate or deposit by absorbing both incident and reflected laser energy. This energy is transferred to the substrate through both heat conduction and radiation heat transfer mecha nisms as well as by vapor condensation with consequent release of latent heat.This process is quite unstable in nature and acts like a positive feedback control system. Usually the reflection of a material decreases markedly upon a solid to liquid phase transition. This usually leads to a corresponding step increase in photon absorption, which in tum leads to a step increase in the temperature rise per unit time. When the vapor pressure equals the system pressure, another step increase in energy coupling to the substrate occurs. This causes the substrate or deposit temperature to rise even faster, greatly increasing the local pressure near the incidence of the laser beam to the substrate.

Analysisof this parameter is totally dependent upon the variables of a specific system.The process of chemical reaction and deposition is complicated and multistep by nature.Bauerle38 has identified seven consecutive steps in laser-induced reactions:1. Transport of reactants into the reaction zone through the boundary layer.2. Adsorption of one or more reactants onto the substrate. 3. Pyrolytic or photolytic activation of molecules near or at the substrate surface. Transport of product atoms or molecules to the substrate surface, with possible recombination or secondary reactions on the way.4.5. Condensation or further reactions of the products of step 3 on the surface.6. Desorption of reaction products from the substrate. 7.Transport of reaction products out of the reaction zone and through the boundary layer.Each of these steps proceeds at a certain rate.

Desorption of reaction products from the substrate. 7.Transport of reaction products out of the reaction zone and through the boundary layer.Each of these steps proceeds at a certain rate. The total process can proceed no faster than the slowest step, the rate limiting step.At low surface temperatures for LCVD or low laser intensities for LECVD the rate of deposition is usually limited by reaction kinetics. Past a critical temperature, the reaction rate becomes diffusion limited. Reactants or reaction products can not diffuse through the boundary layer faster than the reaction proceeds. The reaction rate in this regime is a function of the thickness of the boundary layer, the bulk concentration of the reactants and products, the temperature and the pressure, and the laser pulse time.585112111 Momentum Boundary Layers in LCVD and LECVDA boundary layer develops whenever a fluid flows over a surface with a differentrelative velocity.

A boundary layer is a regime with a different velocity than the bulk fluid. The thickness of a boundary layer is defined as being that distance from the surface where the velocity of the fluid reaches 99 % of the free-stream velocity. The value of the local Reynolds number deter mines the type of the boundary layer:Re ::sxvp Jlxwhere x is defined as the distance from the leading edge of the plate. For the case of flow over a flat plate the following flow regimes are experimentally found:Rex < 2 x 105the boundary layer is laminarthe boundary layer is in transition3 X 106 >Rexthe boundary layer is turbulentHowever, for most real reactor designs, such simplistic analysis techniques have no relevance.Consider an area large relative to the incident laser beam diameter where it is desired to have an even reaction rate so that scan velocities may be constant over the whole area for all scan times.

It is desired then to have an even, non-turbulent boundary layer in order to control the reaction rate if the deposition is in the desired diffusion limited regime necessary for high rate deposition. The best way to determine if such a boundarylayer exists is to perform a finite element analysis of the flow field.5.2112 Diffusion Boundary Layers In LCVD And LECVDreaction reactants and products get to and from the reaction site by both momentum and diffusive transport. When the reaction kinetic rate is sufficiently high, the development of a diffusive boundary layer will result.understand the fundamentals of mass transfer, it is beneficial to review some basicequations.

This results in three dimensional· diffusion. Figure 2c and 2dshow the development diffusion boundary layer. A planer boundary layer will occur.when the scan lines either contacting for LECVD or overlapping for l.,CVD and the time for a complete scan over the total deposition area is short compared to the characteristic diffusion time needed to replenish the depleted zone. This results in one dimensional diffusion identical to large area CVD.The type of diffusion boundary layer in CVD systems is. highly dependent uponThe effect of the type of diffusion boundary layer on the deposition rate is profound. etal.39 claim that for favorable reactant systems, very rapid deposition rates (g.reaterthan 100 JIm) secand scan speeds for line deposition (greater than 10::) can be achieved.

etal.39 claim that for favorable reactant systems, very rapid deposition rates (g.reaterthan 100 JIm) secand scan speeds for line deposition (greater than 10::) can be achieved. However, such depo sition rates are found only for the case of a hemispherical diffusion boundary layer that one would not find occurring in large area rapidly scanned material processing using laser beams.Bauerle40 modeled a simple case assuming negligible surface diffusion and convective flow. He also assumed that the reactant diffusion is much slower than product diffusion. The time dependent surface reaction·f1ux (proportional to deposition rate) for a hemispherical fusion boundary layer is givend (r+~)\Dt2Djet) == If) +d 1+ If) e 0eric[( 1) If) +df) (4Dty~.!oJ}where at time zero the reaction begins.

If) is a length scale proportional to the mean free path the gas and is defined as:I f)== 4DThe time required to reach steady state is:Because Dfor d« if)'tcxp-lfor d» If)'t can be < IOns for small spot sizes and pressures below one atmosphere, such as what occurs in microchemical processing. For steady state deposition, apply L'Hopital's Rule:2DC"" jet -7 00 ) == - - io +dwhich again yields two limiting cases:jet -7 oo)cxPfor d« 10jet -7 oo)cx1 dfor d» 1060[24][26]Copley studied mass transport during LCVD and concluded that reaction rates can be characterized into three regimes depending uPQn the molar ratio of vapor phase reactants to products, bulk mole fractions, and kinetic rates. 41 The slowest rates are kinetically controlled, where no mass transport limits are found for either reaction products or reactants.

41 The slowest rates are kinetically controlled, where no mass transport limits are found for either reaction products or reactants. The second case is where the molar ratio is one or more, where more moles of reactahts are diffusing to the surface than moles of products are diffusing away. The molar average velocity is towards the substrate. This is a diffusion limited regime, and somelimi.tation is placed upon the kinetic rates. The third regime Copley studied is the case when the molar ratio is less than unity, where the outward diffusion of products impedes the inward diffusion of reactants. He calls this case the diffusion plus convective case and the molar average velocity is away from the substrate. This regime provides the most impediment to the kinetic rate.:'i, :;~ ~:useTB,..

This is in direct conflict with the desire to minimize the boundary layer thicknessin order to.increase the deposition rate for LECVD. Themost common method for increasingcleposition rate at constant P is to increase flow velocity over the substrate, thereby minimizing the momentum and diffusion boundary layer thickness. For LCVD systems where the reaction mechanism occurs by surface decom position of reactant, this poses little conflict. However, in LECVD reaction mechanisms, analogies may be drawn to plasma enhanced chemical vapor deposition (PECVD), where610 •• PI... < - - -OasPlow < - - -deposition often occurs by the creation of activated species, usually free or diradicals, in the gas phase and diffusion of these short lived species to the surface where the deposition reaction occurs.

42.43,44,45InLCvn reaction schemes for large area deposition, the spatial confinement is defined by the reaction kinetic and heattransport\equatiQns.for•• sin.~leli11~deposition,the deposit ""ill always have a diameter smaUerthan the beam diameter for endothermic reactions, because the incident beam diameter intensity follows a Gaussian dependence on radius and the reaction kinetics have an exponential dep~ndence on teIllperatur~. For lar~e area deposition, bulk versus surface heating effects.must be investigated. Forexample, considerthe hole in Figure 2c. Assume the dark region around the hole is hotterthan the white hole, thereby allowing for CVD to occur. In order to maintain constant boundary definition with increasing deposit thickness, it is important to insure that the cold area (hole) can transfer heat from its surface as fast as the hot area (dark region) transfers heat into it.

This may be accomplished by transferring heat from the substrate, using a low temperature gas feed, and by decreasing the absorption length to increase the temperature gradient at the surface and minimize bulk heating effects.reaction schemes, the creation of energetic§pecies occurs throughout the incident laser beam diameter. Deposition occurs more evenly throughout the beam incidence and even a little beyond the beam due to diffusion effects. In a flow field, deposition can occur downstream of the beam.InOne measure of the reaction spatial confinement of a plasma is the meanpath of themolecular species. The mean free path of an ideal gas is given by:A:=Kr 121tcrp[26]At 1000 <>K and 1 atm. pressure, A:= 2.34x 10-7 meters for air. The mean free path specifies the average distance gas phase molecules travel between collisions with other gas phase species. One often assumes that a collision of a radical with another gas phase molecule always results in a reaction.

One often assumes that a collision of a radical with another gas phase molecule always results in a reaction. This distance is sufficiently small to ensure adequate spatial confinement. However, if the system pressure is 10-6 atmospheres, this criterion is no longer sufficient. One must resort to more complicated estimates of excited state mean relaxation times of excited species.Design Of The Remote Plasma Solid Freeform Fabrication VaporPhase Reactor System6..1 Design RequirementsHigh rate nucleation and growth of diamond requires a high local partial pressure ofatomic hydrogen. Localized, controlled deposition of diamond requires localized substrate heating apparently within the thermal processing window of 800-1100 <>C. The reaction chamber is designed to adapt on the existing RE:N1PITOFMS system. The system must have mass flow controllers for all of the process gases as well as chamber pressure control.

The reaction chamber is designed to adapt on the existing RE:N1PITOFMS system. The system must have mass flow controllers for all of the process gases as well as chamber pressure control. The substrate heater must heat to 1100 <> C so that large area deposits may be grown without use of the heater must have a temperature controller. The reactor wall must be water cooled to prevent overheating from the substrate heater. The reactor must be designed to that it is convertible so that a plasma may be created in the chamber over the substrate by TMol mode. Parallel gas flow over the substrate surface will lead to thinner boundary layers.626.2 Plasma SystemAtomic hydrogen will be created by a plasma system. A microwave plasma source was chosen over RF due to the greater power coupling available at the shorter wavelength of the microwaves. A downstream microwave plasma system is used, operating at 2.45 GHz and available from ASTEX, Inc. This system is schematically shown in Figure 3.

A downstream microwave plasma system is used, operating at 2.45 GHz and available from ASTEX, Inc. This system is schematically shown in Figure 3. It is capable of supplying up to 1.5 kW of microwave energy. The design of the downstream microwave plasma reactor tube must allow the metastable (neutrals) excited species created in the plasma to be transported to the substrate for low pressure operation in the 10-2Torr range. Therefore the tube wall must be an electrical insulator and withstand temperatures above 1000 ·C in the microwave coupler. The inner tube wall surface is electrically insulating downstream of the plasma coupler to prevent charged specie recombination. A conventional alumina tube (McDanel Type 998) of 1.5" 0.0. is used to transport the metastable specieS to the reaction chamber. For operation at pressures above 100 Torr, a 114" 0.0. alumina tube is used as the reactant supply line in the center of the 1.5" 0.0. tube to increase energy density of the excited stream.

For operation at pressures above 100 Torr, a 114" 0.0. alumina tube is used as the reactant supply line in the center of the 1.5" 0.0. tube to increase energy density of the excited stream. The plasma system is designed so that substrates may be placed in and doWnstream of the plasma so that the effects of downstream recombination on deposit characteristics and growth rate may be studied. The alumina tube may be cooled by a water jacket or heated by a resistance wire wrap as desired.Three Stub TunerMagnetron HeadGas In tRadical StreamCirculator and Dummy LoadFigure 3: The 1.5 KW downstream microwave plasma system.6.3 Reaction ChamberThe reactor chamber is shown in Figure 4 and 5. Table 2 describes the labels in Figure 4. The reactor is water cooled by a double wall jacket. The top lid of the reactor is convertible to allow for a symmetricTMo1mode plasma coupler to fit on it over a 5" ZnSe COzlaser window.

The reactor is water cooled by a double wall jacket. The top lid of the reactor is convertible to allow for a symmetricTMo1mode plasma coupler to fit on it over a 5" ZnSe COzlaser window. In this mode the laser scanning mirrors are placed in the cylindrical coupler. This will allow for a plasma to be created directly over the substrate in the future. Currently, it has a 2" ZnSe CO2 laser window at the top. The laser optics and scanning system are above the reactor. It has 4 ports mounted at 30· from the substrate to allow for videoimaging and in situ laser interferometry of the growing deposit. A capacitive manometer is placed on one of these ports. The reactor has a fast entry port. The bottom lid of the reactor is an ISO~250 flange that is rem.ovable for reactor mounting on the REMPITOFMS system. The substrate heater and hydrocarbon feed line are mounted on the bottom lid. K~type thermocouple feedthroughs are mounted ona port opposite of the door port.

Many materials have these characteristics such as Si and A120 3 • The substrates will be prepared before. each run by a diamond paste polish to below 11lm grit size followed by an ultrasonic rinse for 5 minutes in each offour solvents: glacial acetic acid, reagent grade I water, methanol, and acetone. Before each part is made, an even primary layer of amorphous carbon with a thickness of 25 Ilm is deposited over the entire substrate by sputtering or PECVD. This carbon layer's purpose is to facilitate even nucleation of diamond over the entire substrate wherever the laser is incident. Silicon is transparent to 10.6 Ilm photons below 700 ·C, while amorphous, sp2 bonded carbon is opaque to this photon energy. Although low defect density diamond films are transparent to 10.6 Ilm photon energy, the high rate deposition planned for SFF will result i.n high defect density sp3 bonded deposits that will be opaque.

Deposition rates may be lowered for growth of high quality sp3 bonded films that will have good optical transparency characteristics.6.5 Substrate Holder and 1100 ·C HeaterThe choice of materials and heater design was motivated to achieve several design goals. The heater must be inexpensive. RF induction heating of a graphite substrate was eliminated because of its cost and large heated volume. Halogen lamps were similarly eliminated due their cost and nonuniform heating. A conventional ohmicly heated resistance wire heater was chosen65for economy. The wire is embedded in a ceramic plate mounted on a hydrocarbon feed is directly upstream of the substrate and is designed to evenly distribute the hydrocarbons over the leading width of the substrate heater. The substrate (silicon wafer) will rest on the heater. A greatconcem is the projected Hfetimeof the Cr-AI-Fe alloyh~ating element in<ahydrogenand oxygen atmosphere.

The substrate (silicon wafer) will rest on the heater. A greatconcem is the projected Hfetimeof the Cr-AI-Fe alloyh~ating element in<ahydrogenand oxygen atmosphere. Hydrogenembrittletnent, oxidation, and sulphur induced green rot can both cause heater elementfailure, but the projected lifetime is hardto '-''''''1..&.1.'-''''''''', The heater support and chamber are designed to facilitate heater replacement.support.7 Computational Plasma ModelingMyJti-elementcontainingplasmas are inherently complex systems because ofthe number ()fnon-trivial chemicalreactions that can occur within them. Reproducible and pre the concentra dictable high rate diamond deposition requires a fundamental understanding tions of the suspected diamondvapor phase precursor chemical species within the plasma and downstream of the plasma.

Understanding the kinetic balance between net sp3 bonded carbon deposition and net sp2 bonded carbon deposition also requires an understanding of the con centrations of the relevant vapor phase chemical species.Plasma modeling and plasma diagnostics are necessary to reach this level ofunderstanding, and are absolutely requiredbeforesignificantprocess scale up may be reliably achieved. Carbon and hydrogen containing plasmas have been well modeled in the literature. Kline et al. modeled a methane RF glow discharge plasma using a Monti-Carlo simulation of electron molecular species interactions. 46 The authors concluded that CH38 were the primary carbon deposition species. glow discharge. This assumption is good for species with characteristic time constants below 1 milliseconds, which included most species. The reactive species that exit the glow discharge zone were found to rapidly decay to equilibrium species. Tachibana et al.

The reactive species that exit the glow discharge zone were found to rapidly decay to equilibrium species. Tachibana et al. performed plasma diagnostics and modeled a methane plasma used to deposit amorphous carbon. 47 Their model included 15 neutral chemical species, 20 ionic chemical species and 23 neutral chemical reactions and 33 ionic reactions. Their model predicted that the most abundant neutral radical in the plasma is with respect to time is nearly proportional to thealso used the plug flow model to predict specie concentrationsThey found that the rate of change in the partial pressure ofofforward power of the plasma: IJJ'CH 4MYCH" =-p- CH"They also found that the ratio of1CH" over the RF power in W decreases as the pressure Inc:re'lses. Suzuki et ai. modeled a DC discharge plasma under diamond deposition conditions. 48 diagnosedtheir plasma using the Langmuir probe technique and optical emission spectroscopy.

Suzuki et ai. modeled a DC discharge plasma under diamond deposition conditions. 48 diagnosedtheir plasma using the Langmuir probe technique and optical emission spectroscopy. They modeled their plasma assuming that the neutral species are in a thermal equilibrium state because the degree of ionization is low (-10.7) and the totalpressure is about 200found that almost all of the CH4is decomposed to. and H· at the statistical temperature above4,800 where diamond was observed to grow. They concluded fronl. this result that the VJ.J.JL.l.QL"LVJ.J..of ·C . plays a significant role for diamond growth.The importance of good complete reaction rate data in plasma modeling cannot overemphasized. The highest growth rates ofdiamond deposition from the gas phase are reported to be from oxygen containing vapor phases, but it is not known why. T~e~efore, oxygen.con taining plasmas are worth of study.

OES provides a line of sight averaging of information through the plasma. CARS is not suffi ciently developed at this time to be a useful technique. REMPI provides very good line of sight averaging of information like OES, but resonant frequency overlap among several different species can lead to misinterpretation of experimental data. REMPITOFMS promises to provide good concentration data of the vapor phase transient species immediately above the depositing film down to 10-8 mole fraction. By combining time of flight mass spectroscopy (TOFMS) with REMPI, REMPITOFMS will allow the relative contributions to be made of all species having overlapping resonant ionization frequencies. REMPITOFMS will be used to qualitatively determine transient specie concentrations within and downstream of the plasma. The REMPI TOFMS results will be compared to the computer model results.A 2+1 REMPI process is shown schematically in Figure 6.

The REMPI TOFMS results will be compared to the computer model results.A 2+1 REMPI process is shown schematically in Figure 6. Two photons of the same energy excite an electron in the molecule to an excited state. Another photon of the same energy is sufficient to eject the electron from the molecule. The excited state is short lived and the molecule may also fluoresce, releasing a photon and dropping down to a lower energy level. The exited state may also decay by intramolecular relaxation. Not all REMPI processes are 2+ 1; 3+1 and 4+ 1 ionizations are also found.Figure 7 shows schematically the REMPITOFMS diagnostic station. The remote plasma diamond SFF reactor is placed on top of the REMPITOFMS system, replacing the substrate heater and bottom ISO-250 flange of the reactor with the diagnostic station substrate heater and ISO-250 flange. A molecular beam is introduced into the REMPI column through a hole of 50-250 Jlm in diameter etched in a silicon wafer.

A molecular beam is introduced into the REMPI column through a hole of 50-250 Jlm in diameter etched in a silicon wafer. The sampling volume of the reactor is a hemisphere of diameter 3-5 times the hole diameter. The sampling funnel is designed so that once molecules enter the REMPI column they do not undergo any collisions. The pressure in the SFF diamond reactor is maintained at 0.01-730 Torr by a mechanical pum~ and proportional flow control valve; while the REMPI column is maintained at about 5 x 10- Torr by two tur bomolecular pumps.Specific species are ionized by the tunable dye laser by the REMPI process. Once ionized, they are accelerated and focused by electrostatic lenses to a Chevron plate detector in a con ventional time of flight mass spectrometer. The TOFMS gives all the ions produced by a single laser pulse the same momentum. The velocity of the ions in collision-free flight is a function of the ratio of their charge, q, to their mass, m.

The TOFMS gives all the ions produced by a single laser pulse the same momentum. The velocity of the ions in collision-free flight is a function of the ratio of their charge, q, to their mass, m. TOFMS is well described in the literature. 58REMPITOFMS will answer many questions on the nature of sp3 bonded carbon depo sition. The substrate temperature may effect the concentration of the gas phase species above it. The effect of the distance from the plasma to the substrate on gas phase specie concentrations at various pressures may also be studied. The effect of plasma power, the mole fraction of input67gases t and total flow rates on gas phase specie concentrations may also be studied. Relating the REMPITOFMS results to deposit characteristics and computational models will answer many questions.> CDa.: w z w.. I•2hvEXCITED STATEFLUORESCENCEGROUND STATENONRADIATlveDeCAYFigure 6:Schematic drawing of a 2+ 1 REMPI process.

The SFF diamond reactor is shown on top of the station.9 SummaryMetastable sp3 bonded carbon deposition growth rates have increased from the A/hr range for hot wire techniques in the early 1980' s to 930 ~mlhr for thermal equilibrium argon-hydrogen plasma torches with downstream methane injection in 1989.60Thermodynamic, kinetic, and transport phenomena theories suggesting that even higher growth rates are possible have been discussed. The design of a vapor phase solid freeform fabrication reactor for fabricating sp3 bonded carbon parts has been described. The reactor is designed to allow for high rate deposition of sp3 bonded carbon by both remote plasma LCVD and in situ plasma enhanced LCVD.The problems preventing high rate deposition of sp3 bonded carbon will be attacked through the triad of experimentation, modeling, and diagnostics.

The knowledge gained from these studies will be incorporated into next generation reactors resulting in faster production, greater spatial definition, and more structurally sound solid freeform fabrication parts.6910 AcknowledgementsThe authors would like to acknowledge Steve Lin and Norman Williams for numerous discussions on gas handling and vacuum system design and implementation. The authors greatly appreciatedthe assistance of Justine Hager and Professor Eckerdt of the Chemical Engineering Departnlent for guidance indesignifig RENIl'I1"()FM~compatabilityintothe system design. This project is sponsored by the Texas Advanced Technology Program.70References 1. Bachman, P. and R.

This project is sponsored by the Texas Advanced Technology Program.70References 1. Bachman, P. and R. Messier, "Emerging Technology of Diamond Thin Films," Chemistry and En2ineerin2May 15 198924-39."Changes Produced by High Temperature Treatment of Diamond," in Properties of Diamonc;l, editedNew York: Academic Press, 1979,403-24.J., "History of Diamond Growth at Metastable Conditions," Proceedin2s of the First International . Dim n -Uk Film, edited by Dismukes, J. P., A. J. Purdes,B. S. Meyerson, T. D.m i m Moustakas,. . Spear, V. Ravi, and M. Yoder, The Electrochemical Society, 89-12, 1989, 1-13.n""V"I.V'" W. von, "Ober Die Ausschiedung Von Kohlenstoff In Form A Von Diamant," Zeit$crift FUr ElectroNovember 15, 1911,971-2.5. Angus, J. "History of Diamond Growth at Metastable Conditions," Proceec;lin2s of the First International Symposium on Diamond and Diamond-like Films, edited by Dismukes, J. P.., A. J. Purdes, B. S. Meyerson, T. D. Moustakas, D. E. Spear, F.

P.., A. J. Purdes, B. S. Meyerson, T. D. Moustakas, D. E. Spear, F. V. Ravi, and M. Yoder, The Electrochemical Society, 89-12, 1989, 1-13.6. Company, JuneA . "Synthesis of Diamond Form Carbon Monoxide," Tonawanda Laboratories, Linde Air Products1956.7. Bundy, F.H. T. Hall, H. M. Strong, R. H. Wentorf, "Man-Made Diamonds," ~,176, July 9, 1955, 51-4.8. DeJjaguin, B. V., D. V. Fedoseev, V. M Lukyanovich, B. V. Spitzin, V. A. Ryabov, and A. V. Lavrentyev, "Filamentary Diamond Crystals," Journal of Crystal Qrowth, 2, 1968, 380-4.9. Lander, J. J. andJ. Morrison, "Low Energy Diffraction Study of The {l11} Diamond Surface," Surface Science, 4, 3, May-June 1966, 241-6.10. Angus, John Deposition," 11.10. Angus, John Deposition," 11.Herbert A.and Wayne S. Stanko, "Growth of Diamond Seed Crystals By VaporP. and R. Messier, "Emerging Technology ofDiamond Thin 1989,24-39.Films," Chemistry and En2ineeringB. V. and D. V.

Stanko, "Growth of Diamond Seed Crystals By VaporP. and R. Messier, "Emerging Technology ofDiamond Thin 1989,24-39.Films," Chemistry and En2ineeringB. V. and D. V. Fedoseev, "Epitaxial Synthesis of Diamond In The Metastable Region," RysSian~~.aJ..J:S&Yli~,9,1970,783-8.tsa<;hnlann, P. and R. Messier, "Emerging Technology of Diamond Thin Films," Chemistry and En2ineerin21989 pp. 24-39.14. Johnson, W. L., III, "The Characterization Of Graphitic And Nongraphitic Carbon By LaserRamanMicroprobe Analysis," in Microbeam Analysis edited by A. D. RomigJr. and W. F. Chambers, San Francisco Press, Inc., 1986, 26-8. D. L. Flamm, D. E. Ibbotson, and 1. A. Mucha, "Diamond Crystal Growth By Plasma ChemicalveIDOs:ltlOn," JOyrnal Of AJ)plied PhySics, 63, 5,March 1, 1988, 1744-8.S. and WilliamB. White, "Characterization ofDiamond Films By Raman Spectroscopy," Journal~~:lJ.iU~~;,e, 4, 2, March-April 1989, 385-93.H. S. Kong, and J. T.

and WilliamB. White, "Characterization ofDiamond Films By Raman Spectroscopy," Journal~~:lJ.iU~~;,e, 4, 2, March-April 1989, 385-93.H. S. Kong, and J. T. Glass, "Electron Microscopy of Vapor Phase Deposited Diamond,"~~~~=~~p~p:.b~'&.'n, 5, 4, April 1990, 801-10."Diamond Films Sparkle As They Come To Market," Science, 249, July 6, 1990,27-8.19." The Economist, 316, 7663, July 14, 1990,84.20. Anthony, T. R., W. F. Banholzer, J. F. Fleischer, Lanhua Wei, P. K. Kuo, R. L. Thomas, and R. W. Pryor, "Thermal Diffusivity ofIsotopically Enriched 12C Diamond," Physical ReviewB, 42, 2,July 15, 1990, .1104-1111.21. Suzuki, Kazuhiro, Atsuhito Sawabe, and Tadao Inuzuka, "Characterizations of The DC Discharge Plasma During Chemical Vapor Deposition For Diamond," Applied PhysicS Utters, 53, 19, November 7, 1988, 1818-9.Joshi, Asawari, A. Gangal, and S. K.

Gangal, and S. K. Kulkarni, "Structure and Properties of Diamondlike Carbon Coatings Deposited In RF Plasma From Benzene and Monosubstituted Benzenes," JoyrnaLof ApJ)liedPhysiCS, 64, 12, vecember 15, 1988,6668-72.23. Williams, B. E. and J. T.Glass, "Characterization of Diamond Thin Film~:< Diamond PhaseIqentifjcation, Surface Morphology, And Defect Structures," JournalofMateria1s Research, 4,2, March-April, 1989,373-84.24. Kitahama, Katsuki, Kazuhiko Hirata, Hirohide Nakamatsu,Shichio Kawai, Naoji Fujinlori, Takahiro Imai, ........ ,"'".... Yoshino, and Akira Doi, "Synthesis Of Diamond By Laser-Induced Chemical Vapor Deposition," AJ)plied .!..!.!..,)~~~~, 49, 11, September 15, 1986, 634-5.7125. Kitahama, Katsuki, Kazuhiko Hirata. HirohideNakamatsu, Shichio Kawai, Naoji Fujimori, and Takahiro Imai, "Synthesis Of Diamond By Laser-Induced CVO, in Photon, Beam, and Plasma Stimulated Chemical Processes At Surface,s, MaterialRe$earc;hSgciety Symposia. 75, 1987, 309-15. 26.

33. Pate, B. B., "The Diamond Surface: Atomic and Electronic Structure," Surfjlce Science, 165, 1986,83-142. 34. Joshi, A., "Oxidation Behavior ofCarbon and Diamond Films," in Thermjll Analysis QfMetilllur~icjll Systems, TMS, Warrendale, PA, 1990. 35. Herziger, Dr. Gerd, "The Influence of Laser-Induced Plasma On Laser Materials Processing," in TheIndustrial Laser Annual Hjlndbook, 1986 Edition, edited by David Belforte and Morris Levitt, 1986, pages: 108-15.36. Allen, S. D., J. A. Goldstone, J. P. Stone, and R. Y. Jan, "Transient Nonlinear Laser Heating and Deposition: A Comparison of Theory and Experiment," Journal Of Applied- PhysicS, 59, 5, March 1, 1986, 1653-7. 37. Perry, Robert H. and Don Green, Perry's Chemical En~ineers' HjlndbQQk, Sixth Edition, McGraw-Hill, 1984. 38. Bauerle, Dieter, Chemical Processin~ With Ljlsers, Springer-Verlag, 1986. 39. Allen, S. D., R. Y. Jan, S. M. Mazuk, and S. D.

SIGN OF AN EMPIRICAL PROCESS MODEL AND ALGORITHM FOR THE TUNGSTEN INERT GAS WIRE+ARC ADDITIVE MANUFACTURE OF TI-6AL-4V COMPONENTSFilomeno Martina∗,a, Stewart W. Williams∗, Paul Colegrove∗∗Welding Engineering and Laser Processing Centre (WELPC), Cranﬁeld University, Cranﬁeld, MK43 0AL, United Kingdom af.martina@cranﬁeld.ac.ukAbstractIn the wire+arc additive manufacture process parameters can be varied to achieve a wide range of deposit widths, as well as layer heights. Pulsed Tungsten Inert Gas was chosen as the deposition process. A working envelope was developed, which ensures unfeasible parameters combinations are excluded from the algorithm. Thanks to an extensive use of a statistically designed experiment, it was possible to produce process equations through linear regression, for both wall width and layer height. These equations are extremely useful for automating the process and reducing the buy-to-ﬂy ratio.

These equations are extremely useful for automating the process and reducing the buy-to-ﬂy ratio. For a given layer height process parameters can be selected to achieve the required layer width while maximising productivity.1 IntroductionAdditive manufacturing (AM) is one of the emerging innovative manufacturing approaches. Formerly a rapid prototyping technique, it is been considered as a manufacturing technique in its own right, particularly within aerospace and more recently automotive sectors. Wire+arc additive manufacturing (WAAM) [1] is one of the many AM processes, and it consists in the utilisation of arc welding tools for AM purposes. Different from laser-based processes, which use only one set of deposition parameters and have ﬁxed values for layer height, hatch distance, etc, WAAM offers the possibility of changing the deposition parameters on an ad hoc basis.

Almeida and Williams [1] were the ﬁrst to relate the process parameters to the geometry of the material deposited, for Ti– 6Al–4V, while using an innovative metal inert gas (MIG) approach, namely Cold Metal Transfer (CMT). The model enabled the ﬂexible selection of deposition parameters that changed according to the desired geometry; for instance, to achieve a certain wall width or layer height, a speciﬁc combination of process parameters would be selected. It must be noted that to minimise material waste, being able to deposit a geometry that is as close as possible to the ﬁnal one is highly beneﬁcial.The scope of the present work is to determine a model that relates process parameters to weld bead geometry, for tungsten inert gas (TIG) based WAAM.

Moreover, this work addresses a fea- sibility issue by determining a working envelope for the process, which excludes the unfeasible parameters combinations.2 Experimental methodAn EWM Plasma/TIG power source and an EWM TIG torch mounted on an ABB 6-axes robot were used for the deposition. A 100 mm long trailing shield was used to minimise oxidation by applying argon gas locally. The 1.2 mm wire was always front-fed; a schematic of the setup is shown in Fig. 1. After a preliminary assessment of the signiﬁcance of all the parameters, some of them were kept constant to reduce the number of independent variables investigated, and therefore the complexity of the experiment (see Table. 1).

1). Gas ﬂow rates were also kept constant (Table 2).To determine the process window, a visual, qualitative analysis of 70 preliminary single layer (bead on plate) experiments provided the necessary information for understanding the general be- haviour of the process and determining the limitations of the process parameters, as well as ex- cluding the unfeasible parameters combinations.After establishing the process constraints, a D-Optimal design method [2] was chosen for the experiment, given its capability of taking constraints into account and supporting irregular exper- imental regions. Wire feed speed (WFS) [X1], travel speed (TS) [X2] and current (I) [X3] were selected as the factors in the model. WFS is the speed (m min−1) at which wire is fed into the molten pool (Fig. 1); TS is the travel speed (mm s−1) of the manipulator (and consequently of the torch) along the X axis; and I (A) inﬂuences the heat input.

1); TS is the travel speed (mm s−1) of the manipulator (and consequently of the torch) along the X axis; and I (A) inﬂuences the heat input. 21 straight parts were made of 15 layers each and were 250 mm long. Layer deposition always started from the same end. The part was allowed to cool down to 50 ◦C before depositing a new layer for consistent part geometry.Three-factor second-order polynomial functions were ﬁtted to each of the responses:Y = β0 +3 ∑ i=1βixi +3 ∑ i< jβi jxix j +3 ∑ i=1βiix2 iwhere Y is the predicted response, β0 is the constant process effect, βi is the linear effect of xi, βi j are the interactions of ﬁrst order and βii are the quadratic effects of xi.2.1 MeasurementsThree transverse sections were used for the measurements. During the build process, the de- posited layer height (LH) was measured at 50 mm, 100 mm and 150 mm from the deposition start- ing point.

During the build process, the de- posited layer height (LH) was measured at 50 mm, 100 mm and 150 mm from the deposition start- ing point. The average was used to determine the height increment for deposition of the following layer. For the statistical analysis, given the thermal effect of the base plate, the ﬁrst four layers were analysed one by one (LH1, LH2, LH3 and LH4, as shown in Fig. 2) and for the remaining layers, i.e. from the 5th to the 15th, their overall average was considered:LH =LH5 + LH6 + . ..

. . + LH15 11(1)(2)ClampsZXYGas inletTravel speed (TS)Wire feed speed (WFS)Deposition directionAluminiumTIG torchAttachment to6-axes robotTrailing shieldWire feedercurtainsSubstrateWallFigure 1: Schematic of experimental setup.Table 1: Values for the process parameters which were excluded during the factors selection and therefore kept constant.Electrode-to-workpiece distance Tungsten electrode tip angle Wire diameter ∗ Ip/Ib Pulses duration Frequency ∗Ip and Ib are peak and background currents3.5 mm 30◦ 1.2 mm 1.5 0.05 s 10 HzTable 2: Deposition stages: gas ﬂow rate, and durations.Process gas ﬂow rate (l min−1) Trailing shield gas ﬂow rate (l min−1) Duration (s) aDuration depended on the travel speedPre purge 9 20 10Slope Depositionup 9 20 0.19 20 42 to 125 aSlope down 9 20 0.1Post purge 9 20 20Also the wall widths (WW) were considered separately, according to their dependance on the layer height.

To do this, a series of parallel lines was laid, orientated in the Y direction parallel to the baseplate, on the optical microscope images of the three sections, each line corresponding to a speciﬁc layer height. Consequently, it became possible to infer the shape of the wall for every layer, and a WW could be associated with each of them. As well as in the LH measurements, the ﬁrst four WW were measured separately (WW1, WW2, WW3 and WW4) considering the narrowest point, while the remaining part of the wall was considered as a whole. Furthermore, for the overall WW (equation 3) the average of the measurements of the three transverse sections was considered:WW =WWa +WWb +WWc 3where WWa is the width of the section taken at 50 mm, WWb of the 100 mm one, and WWc of the 150 mm one.

The wall widths of the ﬁrst four layers (WW1 to WW4) used only the central transverse section.WWLH1LH4LH3LH2YZSUBSTRATEWW2WW4WW3WW1Figure 2: Schematic representation of a generic part cross-section. The wall width (WW) takes into account the width of the specimen, at its minimum point, for the ﬁrst four layers separately and then for the remaining part as a whole. The layer height (LH) corresponds to the increase in specimen’s height after the deposition of a new layer. The ﬁrst four layers were considered separately.(3)3 Experimental results and discussion3.1 Working envelopeThe constraints on the parameters are represented by the solid lines in Fig. 3. To ensure addi- tional conﬁdence on process reliability and feasibility, the strength of the constraints b, e, d, f was increased by 10% (dashed lines in Fig. 3).

3. To ensure addi- tional conﬁdence on process reliability and feasibility, the strength of the constraints b, e, d, f was increased by 10% (dashed lines in Fig. 3). Note the current considered in the design of experiment is the average one, given by the average of Ip and Ib described in Section 2.3.2 Process capabilityThe ranges of measured responses are shown in Fig. 4a, and a comparison against competing processes is shown in Fig. 4b. The comparison data, which for the CMT and plasma wire depo- sition (PWD) processes considers only out-of-chamber manufacturing methods, is taken from the work of Almeida and Williams [1], Martina et al. [4], Eze [5], Milewski et al. [6], Tolochko et al. [7]. Note that it was assumed that the wall width equalled the spot size for laser-based processes. Fig. 4a shows how the range of both WW and LH is quite consistent and does not vary depending on the layer number. In Fig.

Fig. 4a shows how the range of both WW and LH is quite consistent and does not vary depending on the layer number. In Fig. 4b, TIG deposition performs similarly to PWD, in terms of both WW and LH, although both are slightly smaller with TIG. The capability of TIG deposition is comple- mentary to CMT, with the latter producing thinner walls with a much greater layer height. This suggests the possibility of combining the two processes to increase the overall capability of a man- ufacturing station. The main advantage of the powder-bed laser deposition process is the small feature size capability, however they are characterised by a much lower deposition rate, which ranges from 100 g h−1 to 200 g h−1 [4, 8].Fig. 4c shows the ranges of wall widths that are achievable for each layer height. The minimum WW (around 5 mm) can be obtained for any LH between 0.6 mm to 1.4 mm. While the low WW34.52002.51.5Wire feed speed (m/min)Current (A)Constraints: wire feed speed vs.

layer height.value stays constant, the high value increases steadily and reaches its maximum at around 14 mm for a LH of 1.5 mm. For LH>1.5 mm the WW range reduces progressively, and eventually for a LH of 1.9 mm the WW ranges from 8 mm to 9 mm only. The process ﬂexibility enables layer height to be ﬁxed for the whole part, while the WW may be varied according to the speciﬁc features that need building.The maximum deposition rate achieved using TIG deposition was 1.2 kg h−1; however, it was constrained by the 4 m min−1 limit of the wire feeder used when doing the preliminary experiments (see Section 2). The time to deposit the 15 layers, considering deposition time only, ranged from 10 min 30 s to 31 min 15 s, and depended on travel speed (cooling time is ignored).3.3 Statistical analysis and process equations calculationTo determine the relationship between the measured response and the statistically signiﬁcantvariables, analysis of variance (ANOVA) was performed.

The responses are:wall widths of the ﬁrst, second, third and fourth layers (respectively WW1, WW2, WW3 and WW4 all in [mm]);wall width of the remaining layers (WW in [mm]); • height of the ﬁrst, second, third and fourth layers (respectively LH1, LH2, LH3 and LH4 all in [mm]);height of the remaining layers (LH in [mm]).3.3.1 Statistical testsTable 3 presents the coefﬁcients of determination (R2) and adjusted-R2 for the models. The R2 coefﬁcients, which according to Montgomery [2] measure the proportion of the variability in the data that can be explained by the model, indicate that over 90% of the variability can be explained by the models. In six cases it is over 95% and in ﬁve over 97%. In practice this should translate to good agreement between expected and actual dimensions. Pred-R2 measures the variability explained by the model for new data and is in reasonable agreement with Adjusted-R2. Finally, the Adeq.

Pred-R2 measures the variability explained by the model for new data and is in reasonable agreement with Adjusted-R2. Finally, the Adeq. Precision tests, which measure the signal-to-noise ratio, are greater than 4, which is the minimum value typically accepted for this term.3.3.2 Analysis of plots3D surfaces and contour plots were generated to represent graphically the regression equations. The WW (Fig. 5) does not seem to depend on the wire feed speed, and it is mostly affected by the(a)(b)Figure 5: Wall width 3D surfaces. (a) wall width as a function of wire feed speed and current (travel speed = 4 mm s−1), and (b) wall width as a function of wire feed speed and travel speed (current = 170 A).(a)(b)Figure 6: Layer height 3D surfaces.

− R2 Adeq. Precision0.9576 0.9228 0.9679 0.9466 0.9435 0.8111 0.83 0.7654 0.7474 0.73847.849 34.901 47.329 35.301 36.139 16.667 20.408 18.736 19.066 16.664P-values<0.0001 <0.0001 <0.0001 <0.0001 <0.0001 <0.0001 <0.0001 <0.0001 <0.0001 <0.0001current and the travel speed. The lack of dependence on the wire feed speed is shown also by its p-value, which is 0.3870. The insensitivity to WFS in terms of WW is the main difference with plasma wire deposition. In fact, Martina et al. [4] showed how in plasma wire deposition an increase in wire feed speed results in an increase in wall width. On the contrary, the effect of travel speed and current is in agreement with what seen previously in plasma wire deposition [4]. Fig. 5b show the effect of the interaction of wire feed speed and travel speed. Also in this case the colouring suggests very little effect of wire feed speed, and the predominance of travel speed.The LH results are plotted in Fig. 6.

Also in this case the colouring suggests very little effect of wire feed speed, and the predominance of travel speed.The LH results are plotted in Fig. 6. Unlike the WW, the wire feed speed affects the response. In the plot of the interaction of the wire feed speed with the travel speed (Fig 6b), the surface shows a more quadratic behaviour given by the effect of the travel speed.The other modelled responses (WW1, WW2, WW3, WW4, LH1, LH2, LH3 and LH4) behavesimilarly to WW and LH, and therefore their speciﬁc plots aren’t be included in this paper.3.3.3 Model validationTo validate the models, two additional walls were built. For Validation 1 the parameters were selected randomly within the working envelope, with WFS = 2.5 m min−1, TS = 4 mm s−1 and I = 150 A. For Validation 2 the parameters were chosen so that a WW of 12 mm could be obtained, with WFS = 2.38 m min−1, TS = 2.57 mm s−1, I = 190 A.

For Validation 2 the parameters were chosen so that a WW of 12 mm could be obtained, with WFS = 2.38 m min−1, TS = 2.57 mm s−1, I = 190 A. The measured values were compared with those predicted by Stat-Ease Design-Expert 7.1 R(cid:13) [3], as shown in Fig. 7.

All values fell within the calculated 95% conﬁdence intervals, except for the LH2 of the ﬁrst specimen (underes- timated), and WW2–WW3 of the second (overestimated).4 ConclusionsIn this research:A working envelope in which the process behaves steadily and delivers acceptable parts has been deﬁned;TIG-based WAAM has been modelled empirically, through a systematic approach (design of experiment), and speciﬁcally D-Optimal design;Modelled responses include: LH, WW for the ﬁrst four layers, and the average for the re- maining ones;The model has been validated successfully.In terms of process capability:Layer height ranged from 0.59 mm to 2.12 mm; • Wall width ranged from 4.93 mm to 13.88 mm; • Maximum deposition rate was 1.2 kg (limited by wire feeder).From a statistical point of view:Wall width is highly dependant on the current; • Layer height is dependant on wire feed speed.AcknowledgementsThe authors would like to thank Mr Flemming Nielsen and Mr Brian Brooks for their support during the research activities.

SIGN OF AUXETIC SANDWICH PANELS FOR STRUCTURALAPPLICATIONSLi Yang1, Ola Harrysson2, Denis Cormier3, Harvey West2, Chun Park2, Kara Peters21University of Louisville, Louisville, KY 2North Carolina State University, Raleigh, NC 27695 3Rochester Institute of Technology, Rochester, NY 14623AbstractBased on an analytical modeling analysis, a sandwich structure with a 3D re-entrant auxetic core was designed. Auxetic samples were produced by electron beam melting (EBM) and selective laser sintering (SLS), and compared to other regular cellular sandwich structures through various experiments. It was shown that sandwich structures with pre-designed auxetic cores could exhibit significantly improved mechanical properties such as bending compliance and energy absorption, which are critical to many structural applications.

This work demonstrated an alternative of effectively designing 3D cellular structures, and also showed the potential of this type of auxetic structure in applications via careful design.IntroductionIn many applications, the sandwich structures are often subject to conflicting requirements such as the need for minimal weight versus the need for high stiffness [1]. Two- dimensional cellular structures and stochastic foams are predominantly used as sandwich cores, but they often lack the level of properties that satisfy all requirements [2-6]. Three-dimensional cellular structures possess many advantages in such applications. However, their practical use is largely hindered by the restrictions of manufacturing processes and the resulting limitation in the development of design theories [6-10].One such three dimensional structure that shows potential is the auxetic structure, or a structure that exhibits negative Poissons’ ratio in one or more directions.

Due to their counterintuitive behavior and exceptional properties, auxetic structures have received considerable attention in past years [11-20]. The combination of high shear modulus and strength [11-12], high toughness [13-16], large bending compliance [17] as well as tailorable elastic modulus and yield strength [18-20] makes auxetic structures especially suitable for sandwich structure applications.Traditionally, auxetic structures have been fabricated via a multi-step process that involves rather complex designs of molds and heat compression processes [11]. Furthermore, the mechanical properties of the resulting structures are largely empirical. Additive manufacturing provides an efficient alternative to fabricate these structures. In combination with design theory, additive manufacturing processes could be implemented to produce functional sandwich parts via design for functionality (DFF).In the current paper, a 3D re-entrant auxetic structure was investigated.

Previous work has established analytical models for some fundamental properties of the re-entrant auxetic structure, and it was shown that the mechanical properties of this type of structure could be tailored to a wide range of different applications [18-22]. Based on this work, this paper demonstrates the design of the re-entrant auxetic structure for sandwich panel applications with the focus of bending compliance and energy absorption abilities. An experimentally based study was carried out to compare two re-entrant auxetic sandwich structures to conventional sandwich structures with non-stochastic cellular cores, and the results demonstrate the feasibility of using auxetic structure for such applications.Structural designsFig. 1(a) shows the unit cell design of the 3D re-entrant auxetic structure. This auxetic structure is an orthotropic structure where directions x and y exhibit identical properties due to the cell’s symmetry.

This auxetic structure is an orthotropic structure where directions x and y exhibit identical properties due to the cell’s symmetry. Therefore, the design of the structure could be represented by the simplified 2D geometry shown in Fig.1(b), which includes the length of the vertical (H) and re-entrant (L) struts, the re-entrant angle θ, and the thickness (t) of the strut (not shown in the figure).a. Unit cellb.

α=H/L, and νzx, νxz, Ez, Ex, σz, σx are the Poisson’s ratios, modulus and yield strength of the re-entrant auxetic structure in the z and x directions, respectively. Note that Eq. (5) and Eq. (6) are in implicit form of the yield strength in the respective directions. Though not shown here, from Eq. (1)-(6) it could be derived that the mechanical properties of the re-entrant auxetic structure have a monotonous relationship with respect to its Poisson’s ratio. In general, the mechanical properties of the re-entrant auxetic structure improve with larger negative Poisson’s ratio. While in agreement with previous experimental observations using stochastic auxetic structures, this relationship also provides an efficient method for quick screening of designs of re-entrant auxetic structures, along with the relative density relationship derived for general cellular structures [23].

After screening the design, a more detailed analysis through either analytical methods or more accurately, finite element analysis, can be carried out to fine tune the designs [19-22].In many applications of sandwich panel structures, the structure is required to withstand large amounts of bending/shearing deflection while maintaining sufficient stiffness. In addition, when such structures are used in aerospace and automobile applications as exterior surfaces, low- energy impact protection from small objects and debris is also an important factor for consideration. Auxetic structures were predicted to possess promising properties in these applications compared to regular cellular structures. Based on the analytical model, it becomes possible to design a 3D auxetic structure that meets both requirements.

Based on the analytical model, it becomes possible to design a 3D auxetic structure that meets both requirements. In this study, the bending performance and low-impact energy absorption ability of the auxetic structures were evaluated in separate sets of experiments.Bending performance Considering the fact that the Poisson’s ratio values can be used as a performance indicator for different designs, two re-entrant auxetic unit cell structures were designed with significantly different Poisson’s ratio values. As shown in Table 1, auxetic structure A1possesses a larger negative Poisson’s ratio value in νzx while auxetic structure A2 possesses larger negative Poisson’s ratio value in νxz.

As shown in Table 1, auxetic structure A1possesses a larger negative Poisson’s ratio value in νzx while auxetic structure A2 possesses larger negative Poisson’s ratio value in νxz. Due to the way the 3D CAD model was created, the thickness of the vertical strut tV differs from that of the re-entrant strut tR for each design.Table 1 Design parameters of auxetic structures for bending performanceDesign H (mm) L (mm)A1 A215 7.5957.5 4θ (Deg.)45 70tV (mm)1 1tR (mm)0.707 0.940νzx1.704 -0.445νxz0.547 -1.659Relative Density 0.063 0.116The sandwich panels for bend testing were designed as illustrated in Fig. 2 using the two auxetic core designs (A1 and A2). The dimensions of the sandwich panel cores were kept at approximately 17 mm x 20 mm x 150 mm while maintaining structural symmetries in each direction. In addition, the thickness of the sandwich skins was also fixed at 1 mm in the design.

In addition, the thickness of the sandwich skins was also fixed at 1 mm in the design. The resulting structures had 1 x 2 x 14 unit cell repetitions for design A1, and 1.5 x 2 x 20 unit cell repetitions for design A2.Figure 2 Sandwich panel with auxetic core for bending evaluationLow energy impact performanceThe samples for low energy impact performance were designed in a similar manner. Two re-entrant auxetic designs with significantly different Poisson’s ratios in all directions, A3 and A4, were constructed using parameters shown in Table 2. The dimensions of the sandwich cores were kept at around 55 x 55 x 30 mm while maintaining structural symmetries in each direction, as shown in Fig.

The dimensions of the sandwich cores were kept at around 55 x 55 x 30 mm while maintaining structural symmetries in each direction, as shown in Fig. 3.Table 2 Design parameters of auxetic structures for low energy impact performanceDesign H (mm) L (mm)A3 A44 7.53.75 3.75θ (Deg.)75 45tV(mm)1 1tR (mm)0.966 0.707νzx0.180 -1.343νxz4.462 -0.610Relative Density 0.211 0.220Figure 3 Sandwich structure with auxetic core for low energy impact performanceExperimentationThe sandwich samples for bending performance were manufactured via the electron beam melting (EBM) process using 100/+325 mesh spherical Ti-6Al-4V powder produced by the plasma rotating electrode process (PREP). Identical default process settings for electron beam melting of lattice geometries were used to produce all samples. All samples were oriented in the build chamber such that the two skin faces were normal to the build direction.

All samples were oriented in the build chamber such that the two skin faces were normal to the build direction. Three samples were made for each type of structure, and the samples were fabricated in two builds.The sandwich samples for the low energy impact performance were produced by Fineline Prototyping Inc. using a 3D Systems SinterStation Pro selective laser sintering (SLS) system with ALM PA650 powder. All samples were built in the same orientation with the face sheets normal to the build direction. Six samples were made for each type of structure. After the samples were made, they were cleaned in the standard powder recycling system for this process.After the samples were cleaned, their dimensions were measured using digital calipers, and their masses were weighed using a digital balance having a resolution of 0.0001g. Tables 3 and 4 show the actual dimensions of the sandwich samples for bending and impact testing respectively.

Tables 3 and 4 show the actual dimensions of the sandwich samples for bending and impact testing respectively. The dimensional accuracy of the samples for both tests was within ±5% compared with the CAD model. It is worth noting that due to the noticeable stair step effect observed with the auxetic samples, the strut sizes of the auxetic cores for A3 and A4 sandwich samples were only about 75% of the designed values.Table 3 Actual sample dimensions for bend testingDesign A1 A2L1 (mm) 20.405±0.259 19.829±0.153L2 (mm) 148.675±0.155 150.334±0.140T (mm) 22.538±0.125 16.222±0.029Rel. Dens. 0.149±0.008 0.214±0.006Table 4 Actual sample dimensions for low energy impact testingDesign A3 A4L1 (mm) 68.449±0.057 53.971±0.054L2 (mm) 68.457±0.025 53.962±0.035T (mm) 26.467±0.051 30.709±0.016Rel. Dens. 0.226±0.006 0.175±0.003The bending performance was evaluated by three-point bend testing as shown in Fig. 4.

Dens. 0.226±0.006 0.175±0.003The bending performance was evaluated by three-point bend testing as shown in Fig. 4. The testing was performed on an Applied Test System 1620C at a constant strain rate of 1.27 mm/min. The span of the support points was 114.3 mm. The support rollers had a diameter of 12.7 mm, and the load roller had a diameter of 25.4 mm.The low energy impact performance was evaluated by the drop weight test as shown in Fig. 5. The testing was performed on a drop weight tester with the same setup used by a different study [24]. The drop impactor and the cross head weighed 5.5 kg in total, and the diameter of the impactor’s hemispherical probe was 19 mm. The samples were clamped between two clamping plates, and the impactor dropped down from a pre-marked height and impacted the samples. The impact energy level used for these tests was between 1-1.3 J (e.g.

The impact energy level used for these tests was between 1-1.3 J (e.g. initial impact velocity 0.6 m/s- 0.7 m/s), and the impact was repeated for up to five times for each sample except for the cases where the top skin of the samples were completely perforated.Figure 4 Three-point bend testFigure 5 Drop weight low energy impact testBend Test ResultsTable 5 shows the results of the three-point bend testing for different structural designs. From Table 5 it can be seen that the two auxetic sandwich structures exhibit significantly different properties. The auxetic sandwich A1 exhibits significantly lower bending stiffness andhigher bending compliance compared with auxetic structure A2. From Table 1, it is known that the auxetic structure A2 has a larger negative Poisson’s ratio νyx, so it is expected to exhibit significantly higher flexural stiffness as well as maximum allowable loading compared to auxetic structure A1.

As a result, it is also expected that sandwich panel A2 would exhibit higher total energy absorption at fracture, which in practice can be interpreted as a measurement of the structural toughness. On the other hand, auxetic sandwich A1 exhibits significantly higher resilience compared to A2 with approximately 50% more maximum deflection. This also contributes to the total energy absorption of the A1 sandwich samples. Furthermore, it was observed that the auxetic sandwich samples did not exhibit any localized stress concentration at the loading point. This could be largely attributed to the negative Poisson’s ratio values of the auxetic cores, which accommodate compressive stress by lateral shrinkage, thereby minimizing the stress concentration and skin wrinkling at the loading point, as can be observed in Fig. 4.Table 5 Bending performance of different structural designsDesign Max. Force (N)A1 A21206.43±37.35 3432.23±58.58Max.

4.Table 5 Bending performance of different structural designsDesign Max. Force (N)A1 A21206.43±37.35 3432.23±58.58Max. Deflection (mm) 9.20±0.77 5.96±0.27Flexural Stiffness (N-m2) 4.10±0.33 17.93±0.80Max. Energy Abs. (J) 8.73±0.50 16.54±1.61Low Energy Impact Test ResultsFig. 6 shows the average total energy absorption and energy absorption per strike for auxetic structures A3 and A4. During the test, each auxetic A4 sandwich sample was perforated after 3 strikes. In the application of low energy impact protection, it is highly desired that the protection structure absorb as much energy as possible at a minimum response force level in order to minimize the impact to the structures being protected. Therefore, the auxetic A4 sandwich exhibited a very desirable combination of high energy absorption and low peak response force that would make it an ideal candidate for such application.

Therefore, the auxetic A4 sandwich exhibited a very desirable combination of high energy absorption and low peak response force that would make it an ideal candidate for such application. Since the auxetic A4 design has a significantly higher negative Poisson’s ratio νzx in the impact direction, it could be expected that the A4 sandwich would exhibit higher strength as well as modulus, which in general indicate higher total energy absorption for cellular structures.It is often the case with additive manufacturing processes that material properties exhibit a certain degree of anisotropy. This can be a significant factor for thin features such as cellular struts. It was expected that the sandwich samples were weakened as a result of the specific build direction. Tensile tests were therefore performed on coupons fabricated in three orientations (0°, 45°, and 90°) with respect to the vertical build axis in order to provide an estimation of the anisotropy of the parts. Fig.

Fig. 8 shows the relationships of the mechanical properties of the structures based on the build orientation. Through simple analysis it is known that the auxetic A4 sandwich sample struts had the smallest build orientation angle (i.e. the angle between the strut axis and the build direction). Therefore, it could be expected that the A4 samples were further weakened by another 20-30% due to the strut orientation. When considering the fact that the auxetic samples were weakened due to the reduced core strut size, it is reasonable to concludethat the auxetic A4 design with the large negative Poisson’s ratio would result in significantly improved low energy impact protection.As a conclusion, it could be seen that through efficient design screening, the auxetic sandwich structures could be designed to meet the specific needs of different applications.

In the case of applications with multiple objectives, an optimization method could be implemented to achieve a good balance between design extremes. For example, in this study, it is shown that in general, the bending performance and the impact energy absorption are two conflicting requirements for auxetic sandwich structures.

For example, in this study, it is shown that in general, the bending performance and the impact energy absorption are two conflicting requirements for auxetic sandwich structures. Therefore in order to design structures that meet both requirements, weighing factors will need to be chosen, and a compromise solution will be achieved to satisfy design requirements.4.54.21Total energy absorption4) J ( n o i t p r o s b A y g r e n E3.532.521.510.84Energy absorption per strike3.111.040.50A3A4Figure 6 Total energy absorption and average energy absorption for various sandwich designs)N( d a o L k a e P2000 1800 1600 1400 1200 1000 800 600 400 200 017571102A3A4Figure 7 Peak response force for various sandwich designs(a) Modulus Figure 8 Anisotropy of the parts made in SLS(b) StrengthConclusionIn this study the bending performance and low energy impact performance of auxetic sandwich structures were assessed.

It was observed that through design the auxetic sandwich structures could exhibit significantly varied performance in each application cases. Further work that applies optimization to the analytical model for multi-objective designs could be done on the basis of the current work, and could potentially make significant contribution to the development of design for additive manufacturing (DFAM) theories.Reference[1] H.G. Allen. Analysis and design of structural sandwich panels. Pergamon Press, Oxford, UK. 1969.[2] T.M. McCormack, R. Miller, O. Kesler, L.J. Gibson. Failure of sandwich beams with metallic foam cores. International Journal of Solids and Structures. 38(2001): 4901-4920.[3] Gaetano G. Galletti, Christine Vinquist, Omar S. Es-Said. Engineering Failure Analysis. 15(2008): 555-562.[4] V. Crupi, R. Montanini. Aluminum foam sandwiches collapse mods under static and dynamic three-point bending. International Journal of Impact Engineering. 34(2007): 509-521.[5] A.

sign of Embedded Resistance Heating Element using Rapid Manufacturing Processsubjected to prolonged heating and the ﬁnal top layer will have a very rapid cooling rate. This varies the micro structure through out the part.The conditions in the laser deposition process can be compared to the traditional heat treatment process. Prior to deposition the substrate is at room temperature. The 1st layer of deposit undergoes immediate quenching upon solidiﬁcation. High alloyed steel, such as H13 tool steel when heated above the austenitizing temperature partially dissolves the carbide and alters the matrix. The ferrite is transformed to austenite. If this steel is quenched rapidly in the hardening process, the carbon atoms do not have time to reposition themselves to allow the reforming of ferrite from austenite. Instead they are ﬁxed in a position where they really do not have enough room. The result is high microstresses that can be detected as increased hardness.

Instead they are ﬁxed in a position where they really do not have enough room. The result is high microstresses that can be detected as increased hardness. This hard structure is called martensite. When H13 tool steel is hardened the matrix is not completely converted into martensite. Some austenite is always left, known as ”retained austenite”. Hence, after cooling, the steel has a microstructure consisting of martensite, carbides, and retained austenite. This structure has high inherent stresses. This can cause cracking. Preheating the substrate avoids immediate quenching of the laser deposit. H13 tool steel deposit can be gradually cooled down if the substrate is preheated. This avoids the high inherent stresses in the deposited part.In the traditional heat treatment process material should be tempered immediately after quenching.

This avoids the high inherent stresses in the deposited part.In the traditional heat treatment process material should be tempered immediately after quenching. It is also to be made sure that quenching should be stopped at a temperature of 50-70◦C and tempering should be done at once.Laser deposition is a layer by layer process. The 1st layer deposited on the substrate is subjected high cooling (quenching). Next layer is deposited and the 1st layer is reheated. This is an analogy to the traditional heat treatment process. As mentioned above, to obtain hardness, toughness and dimensional quality, the quenching should be stopped at particular temperature (usually 50-70 ◦C) [1].From the above few studies it is seen that preheating the substrate is necessary to obtain good qualityparts without cracks.2 Preheating of SubstratesLaser welding of tool steel - The main problem in welding tool steel is its high hardenability.

Weld cools quickly once the heat source is removed and the weld metal and part of the heat aﬀected zone will harden. This transformation causes stresses with concomitant risk of cracking. Due to this fact tool steel cannot be welded at room temperature without considerable risk of cracking. Hence preheating becomes a necessary criterion for laser welding [2].Laser Deposition - Laser deposition is very much similar to laser welding process. As mentioned in the introduction, laser deposition process is an analogy to traditional heat treatment process and in order to obtain good quality parts it is necessary to maintain a particular temperature during quenching. This is where preheating becomes a necessity.Machining - The main factor that aﬀects machining is the chatter. The amount of chatter during ma- chining will very much be reduced when the part is preheated [3]. So this makes machining easy and further more high surface ﬁnish can be achieved.

The amount of chatter during ma- chining will very much be reduced when the part is preheated [3]. So this makes machining easy and further more high surface ﬁnish can be achieved. This is another application of the preheating the substrate.2.1 Advantages of PreheatingReduce the risk of hydrogen cracking: Cracking is caused due to the diﬀusion of hydrogen to the highly stressed parts of the deposition.Reduce the hardness of the weld heat aﬀected zoneReduce shrinkage stresses during cooling and improve the distribution of residual stresses.Reduces the chatter during machining [3]Eighteenth Annual Solid Freeform Fabrication Symposium2Design of Embedded Resistance Heating Element using Rapid Manufacturing ProcessResidual stresses can be inhibited to a great extend by preheating. This in turn prevents hydrogen cracking.

This in turn prevents hydrogen cracking. Hence from the above mentioned reviews it is found that preheating the substrate is a necessity for not only laser deposition process but also laser welding process too.3 The Preheating elementThe previous sections elaborate the need for a preheating element to preheat the substrate and in this section the procedure followed to design and fabricate the preheating element is brieﬂy explained.The basic idea of the preheating element is to manufacture a substrate with an embedded resistance heating element it. This substrate will then be used as the base to deposit parts. The method of fabrication of this preheating element is mentioned in the following paragraphs.The process used for the fabrication of the preheating element is the laser aided manufacturing process (LAMP), developed at the University of Missouri - Rolla [5–13].

It is done with the help of a 3016L Fadal CNC machine which is integrated with a 1 KW diode laser with a ﬂat power distribution. The Fadal 5 axis CNC machine used in LAMP lab consists of an X,Y table, a movable Z axis and an A and B rotary axis.A heating element is fabricated via the following procedure:1. A 4x4 inch substrate which is of tool steel is ﬁxed on the vice.2. H13 tool steel powder is to be deposited on this substrate.3. A 2x3 inch pattern as required is deposited on the substrate.The deposit pattern is shown below. As shown ﬁgure 2 the channels in between the deposit is for the placement of a resistance heating element which heats up the substrate on connecting it to a power supply.Figure 2: Heating element deposition path. The red line indicates the intended position of the Ni-Cr element.At ﬁrst the pattern is deposited using direct metal deposition process.

The red line indicates the intended position of the Ni-Cr element.At ﬁrst the pattern is deposited using direct metal deposition process. A resistance heating element, which in this case is Ni-Cr wire is cast in the pattern using a castable refractory material, Ceram Cast.Eighteenth Annual Solid Freeform Fabrication Symposium3Design of Embedded Resistance Heating Element using Rapid Manufacturing ProcessAfter casting, the top portion is attached via laser welding. In order to make sure that the welding takes place correctly, the complete deposition, machining and welding processes should be done one after the other without removing the substrate from the vice.Figure 3: Heating element structure with castable refractory material added.Conduction of heat should take place through the top welded plate and not through the side.

If the heat is dissipated through the sides then eﬃcient heating will not take place and this will not ensure the proper use of the preheating element.In the design of the Heating element the primary importance is the conduction of heat. Heat should be conducted through the top welded portion. In order to achieve this, the channel was separated from the outer wall by air gap Figure 4. This would ensure less amount of conduction of heat through the sides.Unfortunately the air gap does not provide any support to the top weld plate during machining. In order to overcome this issue triangular structural supports are deposited in between the air gaps. This provides suﬃcient support to the weld plate during machinng.Figure 4: CAD representation of the deposited heating element structure.The pattern is deposited and the resistance heating element is cast in the pattern using ceram cast. Ceramcast provides electrical insulation.

Ceramcast provides electrical insulation. The resistance heating element is connected to a DC power supply.Eighteenth Annual Solid Freeform Fabrication Symposium4Design of Embedded Resistance Heating Element using Rapid Manufacturing Process4 Future WorkFabrication of the heating element using laser aided manufacturing processTesting the Microstructural eﬀect.Comparison of the Microstructure with and without the use of preheating element.Provide a constant temperature during deposition process to change the microstructure.References[1] Uddeholm Tooling AB. Heat treatment of tool steel, 2000.[2] Uddeholm Tooling AB. Welding of tool steel, 2000.[3] A. K. M. N. Amin and M. Abdelgadir. The eﬀect of preheating of work material on chatter during end milling of medium carbon steel performed on a vertical machining center (vmc). Transactions of the ASME, 125:674–680, November 2003.[4] S. Ghosh and J. Choi.

In Proceedings of the Twelfth Annual Solid Freeform Fabrication Symposium, page 138, August 6-8 2001.[8] Frank W. Liou, Jianzhong Ruan, Heng Pan, Lijun Han, and M.R. Boddu. A multi-axis hybrid man- ufacturing process. In Proceedings of the 2004 NSF Design and Manufacturing Grantees Conference, 2004.[9] Jianzhong Ruan, Kunnayut Eiamsa-ard, Jun Zhang, and Frank W. Liou. Automatic process planningof a multi-axis hybrid manufacturing system. In DETC, September 29 - October 2 2002.[10] Jianzhong Ruan and Frank W. Liou. Automatic toolpath generation for multi-axis surface machining in a hybrid manufacturing systemg. In Proceedings of the 2003 ASME Design Automation Conference, Chicago, Illinois, September 2-6 2003. Paper No. DAC-48780.[11] Jianzhong Ruan, Jun Zhang, and Frank W. Liou. Support structures extraction for hybrid layeredmanufacturing. In DETC, 2001.[12] Todd Sparks, Heng Pan, and Frank W. Liou.

Design of Experiments Approach Defining the Relationships Between Processing and Microstructure for Ti-6Al-4VT.A. Wallace, K.S. Bey, K.M.B. Taminger, R.A. HafleyNASA Langley Research Center, Hampton, VA 23681AbstractA study was conducted to evaluate the relative significance of input parameters on Ti-6Al-4V deposits produced by an electron beam freeform fabrication process under development at the NASA Langley Research Center. Five input parameters where chosen (beam voltage, beam current, translation speed, wire feed rate, and beam focus), and a design of experiments (DOE) approach was used to develop a set of 16 experiments to evaluate the relative importance of these parameters on the resulting deposits. Both single-bead and multi-bead stacks were fabricated using 16 combinations, and the resulting heights and widths of the stack deposits were measured.

Both single-bead and multi-bead stacks were fabricated using 16 combinations, and the resulting heights and widths of the stack deposits were measured. The resulting microstructures were also characterized to determine the impact of these parameters on the size of the melt pool and heat affected zone. The relative importance of each input parameter on the height and width of the multi-bead stacks will be discussed..IntroductionElectron beam freeform fabrication, EBF3, is a promising new process for directfabrication of metal components. There is a growing interest in industry for applying this technology for adding details onto simplified preforms, fabricating complex structures, and repairing worn or damaged parts. [1] However, process repeatability is paramount to insertion of this technology from the research environment into industrial applications.

[1] However, process repeatability is paramount to insertion of this technology from the research environment into industrial applications. A thorough understanding of the process is necessary to develop closed-loop process control to attain consistency in the parts fabricated. [2,3] In this study, a design of experiments (DOE) was conducted [4] as a first step in understanding the EBF3 process so that full computerized process control may be achieved. DOE is a systematic approach to minimizing the number of experiments required to obtain the breadth of the processing envelope and the influence of the process parameters on the geometry, dimensions, and quality of the deposited part.The EBF3 process uses a focused electron beam in a 10-5 Torr vacuum environment tocreate a molten pool on a metallic substrate. The part is translated with respect to the beam while wire is fed into the molten pool to build up the part in a layer-additive fashion.

The part is translated with respect to the beam while wire is fed into the molten pool to build up the part in a layer-additive fashion. The tool path is controlled by G-code (a commonly used numerically-controlled machining code) that is generated either by directly programming the EBF3 system or by post-processing a Computer- aided design (CAD) drawing of the part.104Experimental ProceduresDesign of ExperimentsAn experimental test schedule was selected using a DOE approach to examine the role of various process parameters on the resulting EBF3 deposits. This streamlined approach minimizes the number of experiments required and provides a systematic way to evaluate the influence of individual parameters and their interactions on the output. A two-level design, where each parameter is assigned a high or low value, was used in this study.

A two-level design, where each parameter is assigned a high or low value, was used in this study. A set of unique experimental conditions are generated and the appropriate experiments performed by systematically varying the combinations of high and low parameters. A suitably-constructed “factors” array, formed by normalizing the parameter’s high and low values to +1 and -1, serves as a basis for statistical analysis of the experimental output. The results of the analysis are used to evaluate the relative importance of the parameters and their interactions and to generate empirical process models that can be used to guide subsequent production runs.Five input parameters and two outputs were chosen in order to produce a manageablenumber of experiments. The input parameters chosen were beam voltage (V), beam current (C), translation speed (T), wire feed rate (W), and beam focus (F).

The input parameters chosen were beam voltage (V), beam current (C), translation speed (T), wire feed rate (W), and beam focus (F). Although there are many other input parameters that affect the EBF3 process, these were selected because they are the most easily adjusted and believed to be the most influential on the process. The high and low values for the focus were interpreted as a defocused and focused beam, respectively. A half-factorial DOE for five parameters at 2 levels results in the 16 experiments described in Table 1. The interaction of beam focus with the remaining four parameters and four-way interactions were not evaluated to simplify the analysis.

The interaction of beam focus with the remaining four parameters and four-way interactions were not evaluated to simplify the analysis. This allowed evaluation of the five parameters plus all two- and three-way linear interactions among voltage, current, translation speed, and wire feed rate with the16 experiments.EBF3 ProcessingExperiments in this study were performed using Ti-6Al-4V base plate that was 6 inches wide by 12 inches long by 0.2 inches thick and Ti-6-4 wire that was 0.063 inches diameter. The Ti-6-4 base plate was fixtured to the positioning table in the EBF3 system. Three type K thermocouples were attached to the back of the base plate for temperature measurements. The thermocouples were evenly-spaced across the width of the plate and located in the center along the length. One panel was produced for each of the processing parameter combinations in the DOE, consisting of three 10-inch long parallel lines deposited on a single base plate.

One panel was produced for each of the processing parameter combinations in the DOE, consisting of three 10-inch long parallel lines deposited on a single base plate. The first pass for each line deposited was performed at the same translation speed as the experiment to be performed, but at half power, to remove oxides and preheat the base plate for improved bonding of the deposit. The first line consisted simply of this preheat pass. The second line was the preheat pass followed by a single bead deposit. The third line was a preheat pass followed by approximately eleven beads deposited in rapid succession, which is termed a stack.105Table 1.

The third line was a preheat pass followed by approximately eleven beads deposited in rapid succession, which is termed a stack.105Table 1. Test schedule of EBF3 process parameters resulting from the DOE approachExperiment Number 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 Notes:EBF3 Process Parameters Voltage1 Current1 Translation Speed1 Wire Feed Rate1 Focus21 -1 +1 +1 -1 -1 +1 +1 -1 -1 +1 +1 -1 -1 +1 +1 1 High: +1, Low: -1 2 Defocused beam: +1, Focused beam: -11 -1 -1 -1 -1 -1 -1 -1 +1 +1 +1 +1 +1 +1 +1 +11 -1 -1 -1 +1 +1 +1 +1 -1 -1 -1 -1 +1 +1 +1 +11 +1 -1 +1 -1 +1 -1 +1 -1 +1 -1 +1 -1 +1 -1 +11 +1 +1 -1 +1 -1 -1 +1 +1 -1 -1 +1 -1 +1 +1 -1The widths and heights of the stacks were measured using calipers. Heightmeasurements were determined by measuring the total height and subtracting the average base plate thickness.Metallurgical specimens were also removed from selected panels and prepared for metallurgical analysis.

Samples were prepared using standard metallographic preparation techniques and imaged using bright field in an optical microscope. This allowed verification of the height and width measurements and comparison of the heat affected zone and melt pool depths using scaled micrographs.Results & DiscussionExperimental ResultsThe caliper measurements of the stack widths were found to correlate well with selectedstacks measurements using metallographic techniques. Height measurements were more problematic. Many of the experimental conditions produced “bulging” on the back side of the base plate which made measurement with calipers difficult. This effect was encountered on all experiments produced with the low translation speed, presumably caused by the increased heating of the panel as the beam takes more time to make its traverse. Selected specimens were mounted and polished, and the height of the bulge was measured from scaled micrographs.

Selected specimens were mounted and polished, and the height of the bulge was measured from scaled micrographs. To106correct the caliper measurements for this bulge the average value of the bulge height was then subtracted from the caliper measurements for panels which were deformed.deposit1Exp T W -1 1 -1 +1 -1 3HAZ0.8bulge5 mm0.6exp. 1Experiment 1Experiment 3θ0.4exp. 3Figure 1. Micrographs of multi-bead stacks produced by experiments 1 and 3.0.2005001000time (sec)Figure 1 illustrates this bulge as well as some of the variation in width and height that was encountered. This figure compares a multi-bead stack produced by experiment 1 which produced a bulge with experiment 3 which did not. These two deposits were made with the same voltage, current, and wire feed rate, but experiment 1 was produced with the low translation speed and experiment 3 the high.

These two deposits were made with the same voltage, current, and wire feed rate, but experiment 1 was produced with the low translation speed and experiment 3 the high. The figure shows that experiment 3 produced a narrow stack with a small heat-affected zone, while experiment 1 produced a wider stack, a heat-affected zone (HAZ) that extends deep through the thickness, and distortion on the backside of the base plate. The increased width of the stack shown produced by experiment 1 is consistent with the low translation speed, as this allows for a greater volume of material to be deposited per length of deposit. Also, a deeper heat affected zone is consistent with higher heating from the low translation speed.

Also, a deeper heat affected zone is consistent with higher heating from the low translation speed. The higher heating of the lower translation speed is also evidenced in the thermocouple data (normalized to the maximum measured temperature) presented in Figure 2, showing that experiment 1 produced higher back side temperature compared to temperatures compared to experiment 3.Figure 2. Normalized temperature history on the back side of the plate during the multi- bead stacks produced by experiments 1 and 3The normalized data for the thermocouples located directly below the deposit on the backside of the plate for single bead runs are shown in Figure 3. Each line on the graph represents a particular set of processing parameters, in this case those for experiments 1 through 4. For each line, there is an initial peak in temperature that corresponds to the preheat pass at half power followed by a higher peak that corresponds to the deposit of a single bead.

For each line, there is an initial peak in temperature that corresponds to the preheat pass at half power followed by a higher peak that corresponds to the deposit of a single bead. These results again show that the low translation speed (experiments 1 and 2) produced a higher back side temperature than the high translation speed (experiments 3 and 4).1070.8θ0.60.4exp. 1exp. T W F -1 1 -1 -1 -1 +1 +1 2 +1 -1 +1 3 +1 +1 -1 4exp. 20.2exp. 4exp. 30050100time (sec)150200Figure 3- Normalized temperature history on the back side of the plate during thesingle bead runs of experiments 1-4.Micrographs of the melted region produced by the pre-heating pass in experiments 3 and4 are shown in Figure 4 to compare the effect of beam focus. These two experiments were performed at the same voltage, current, and translation speed, and since it was the pre-heating pass, the wire feed rate can be ignored.

These two experiments were performed at the same voltage, current, and translation speed, and since it was the pre-heating pass, the wire feed rate can be ignored. The normalized data from the thermocouple located directly below the beam on the back side of the plate for the heating pass of experiment numbers 3 and 4 is shown in Figure 5. Beam focus is the only processing parameter that differs for these runs. The sharp focused beam results in a higher backside temperature than the defocused beam, but both peak temperatures are low. The micrographs in Figure 4 show the melted region and heat affected zone for these passes to be similar in size and shape. Figure 6 shows higher magnification micrographs for the samples presented in Figure 4. Again, although there is arguably a small difference between these two specimens, the two experimental conditions produced microstructures that are remarkably similar in size and morphology.

Again, although there is arguably a small difference between these two specimens, the two experimental conditions produced microstructures that are remarkably similar in size and morphology. These results suggest that changing the beam focus produces a negligible effect on the width of the melted pool and can, therefore, be discounted as an important parameter.0.21 mm1 mm0.15Exp. 3 defocused beamExp. 4focused beamθ0.1exp. 3: defocused beamFigure 4 – Micrographs of resultsproduced by the pre-heating pass of experiments 3 and 4.0.05exp. 4: sharp focus00200time (sec)400600108Figure 5 – Micrographs of resultsproduced by the pre-heatingExp. 30.5 mmExp. 40.5 mmFigure 6.

4: sharp focus00200time (sec)400600108Figure 5 – Micrographs of resultsproduced by the pre-heatingExp. 30.5 mmExp. 40.5 mmFigure 6. Higher magnification micrographs of pre-heating passes performed for experiments 3and 4.Figure 7 compares micrographs of the microstructure produced by the single-bead stack produced by experiments 1 through 4 (See Figure 3 for corresponding thermocouple data.) All four were produced at the same voltage and current setting, and if the focus can be discounted by the previous argument, then these micrographs show the differences produced by varying translation speed and wire feed rate. For instance, comparing experiments 1 and 3, and similarly experiments 2 and 4, shows the differences produced by translation speed. Experiments 1 and 3 were both produced with the low wire feed rate, and experiments 2 and 4 with the high wire feed rate, but with differing translation speeds.

Experiments 1 and 3 were both produced with the low wire feed rate, and experiments 2 and 4 with the high wire feed rate, but with differing translation speeds. Both of the comparisons show that the faster translation speed produces a bead with smaller width, height, and depth of heat affected zone. Comparing experiments 1 and 2, and similarly experiments 3 and 4, shows the differences produced by varying wire feed rate. Experiments 1 and 2 were both produced with the low translation speed, and experiments 3 and 4 with the high translation speed, but with differing wire feed rates. Comparison of the micrographs show that increasing the wire feed rate leads to an increase in deposit height but no measurable change in width. The change in height is less noticeable for the fast translation speed (experiments 3 and 4), but this would be expected due to the smaller volume of material being deposited.109Exp. 15 mmExp. 2exp. T W -1 -1 1 -1 +1 2 +1 -1 3 +1 +1 4Exp. 3Exp.

15 mmExp. 2exp. T W -1 -1 1 -1 +1 2 +1 -1 3 +1 +1 4Exp. 3Exp. 4Figure 7. Micrographs of single bead passes produced by experiments 1 through 4.DOE Analysis ResultsOne panel per condition has been completed to date, and the resulting stack height andwidths were measured. This lack of replication does not allow for a full analysis of the data and generation of a meaningful process model. However, the results were sufficient to examine the relative importance of the parameters by way of marginal mean plots and Pareto diagrams of the half effect values. [4]These analyses start with averages of appropriate sets of the output of interest, in this case height or width. For each input parameter, j, and therefore each column of Table 1, there are two marginal means: an average of the output corresponding to a high parameter value (+1), denoted -.

The marginal mean plots provide a visual way to evaluate+- mj|/2. Again, the magnitude of theThe marginal mean plots for stack widths are shown in Figure 8 for the input parameters(V = voltage, C = current, T = translation speed, W = wire feed rate, and F =focus), and the corresponding half effect chart is shown in Figure 9. These results suggest that the most significant parameter affecting the width of the deposit is translation speed, with decreasing effect for voltage, current, and wire feed rate. Focus appears to have the smallest effects on the stack width. The marginal mean plot also shows how the stack width varies with each parameter; increasing voltage or current produces an increase in stack width, while increasing translation speed or wire feed rate produces a decrease in stack width. The result of extending the half effects analysis to include interactions between the parameters is shown in Figure 10.

The result of extending the half effects analysis to include interactions between the parameters is shown in Figure 10. These110results show that none of the interactions between parameters appear to have a significant impact on the resulting stack width. However, interactions between current, translation speed, and wire feed rate; translation speed and wire feed rate; and voltage and wire feed rate are more significant than beam focus.0.40.40.40.40.4ni,h t d W k c a ti0.30.20.30.20.30.20.30.20.30.2S e g a r e v A0.10.10.10.10.101 +101 +101 +101 +101 +1VoltageCurrentTranslation SpeedWire FeedFocusFigure 8. Marginal mean plots for average stack widthsh t d W k c a t S r o fi0.10.05s t c e f fEf l a H0TVCWFFigure 9.

Marginal mean plots for average stack widthsh t d W k c a t S r o fi0.10.05s t c e f fEf l a H0TVCWFFigure 9. Stack width half-effects for input parameters with no interactions.111h t d W k c a t S r o f s t c e f fi0.10.05Ef l a H0TC T WT WV WFC TV C TV TV CV CCV T WFigure 10. Stack width half-effects for the input parameters with some interactions.The marginal mean and half effects analysis results for the stack height of 11 beads are shown in Figures 11 and 12. The data for stack height is not as complete as the data for stack width for two reasons.

The data for stack height is not as complete as the data for stack width for two reasons. First, the panels produced at a low translation speed bulged at the back side of the multi-bead stacks making accurate height measurements difficult; the present analysis used the heights after correction for this bulge (See earlier discussion.) Second, due to processing difficulties, several of the stacks were produced with fewer beads, and the height measurements could, therefore, not be included in this analysis. Further work will be necessary to conduct a more thorough analysis; however, the results in Figures 11 and 12 still show some useful general trends. As in the analysis of stack width; translation speed, voltage and current are all significant parameters. In this case, however; increasing the translation speed, voltage, or current will cause a decrease in the stack height. The wire feed rate is the parameter with the most impact on the resulting stack height.

The wire feed rate is the parameter with the most impact on the resulting stack height. Focus still has a small effect, but the results suggest that it has a larger effect on the height than it has on the width.The results of the DOE analysis are all intuitively consistent with the EBF3 process.Several experimental parameters considered in this study would influence the deposit width by controlling the amount of input energy. Anything that would increase the input energy would be expected to increase the size of the melt pool, thus increasing the width of the deposit. Current and voltage would have an obvious impact on the input energy, and increasing either parameter would be expected to increase the deposit width. However, the translation speed of the beam would also influence the total input energy. Decreasing this speed would increase the amount of time the beam spends over a particular location, increasing the energy input and thus increasing the deposit width.

Decreasing this speed would increase the amount of time the beam spends over a particular location, increasing the energy input and thus increasing the deposit width. These arguments are consistent with the DOE analysis of stack width (Figure 8) that showed width increased with increased current or voltage or decreased translation speed..1120.40.40.40.40.4ni, t h g i e H k c a t S e g a r e v A0.30.20.10.30.20.10.30.20.10.30.20.10.30.20.1000001 +11 +11 +11 +11 +1VoltageCurrentTranslation SpeedWire FeedFocusFigure 11. Marginal mean plots for average stack heightst h g e H k c a ti0.1S r o f0.05s t c e f fEf la H0TVCWFFigure 12. Stack height half-effects for input parameters with no interactions.Stack height would also be expected to vary with voltage, current, or translation speed bythe same argument.

Stack height half-effects for input parameters with no interactions.Stack height would also be expected to vary with voltage, current, or translation speed bythe same argument. As the deposit width increases with voltage or current, the resulting height of the deposit must decrease as the total volume of material being deposited stays constant. Translation speed would have this same effect, but in this case the volume of material being deposited also changes. Reducing the translation speed increases the volume of material, countering the decrease in height produced by the higher energy. The DOE analysis results shown in Figure 11 show that decreasing the translation speed increases the deposit height, indicating that for these experimental conditions, the increase in volume is high enough to overcome the decrease in height due to increased energy input.

Further study of this output would be useful, including examination of the parameter inputs.113The effect of wire feed rate is a little more difficult to understand. Clearly, increasing thewire feed rate would be expected to increase the volume of material deposited and, therefore, would be expected to increase height, which is consistent with the results presented in Figure 11. However, this higher volume of input material will require more input energy to melt and less energy will be available to sustain the molten pool. Thus the molten pool diameter will be reduced resulting in a decreased deposit width. Therefore, at a constant heat input (voltage, current, and translation speed held constant), the width of the deposit would be expected to decrease with an increase in wire feed rate.

Therefore, at a constant heat input (voltage, current, and translation speed held constant), the width of the deposit would be expected to decrease with an increase in wire feed rate. This is consistent with the DOE results shown in Figure 8, and is also consistent with observations of EBF3 deposited Al [6].Future WorkMore data must be generated in order to conduct a more complete DOE analysis, andgenerate a meaningful process model. This includes producing replicate deposits for each set of experimental parameters. Extension of the analysis to include other output parameters, such as melt pool depths and heat affected zones, would also be useful.

Extension of the analysis to include other output parameters, such as melt pool depths and heat affected zones, would also be useful. In addition, further work is required to determine the effect of the deposit parameters on the resulting microstructure of the Ti-6Al-4V deposits.Concluding RemarksA study was conducted to evaluate the relative significance of input parameters ondeposits produced by an electron beam freeform fabrication process under development at the NASA Langley Research Center. Five parameters where chosen for study (beam voltage, beam current, translation speed, wire feed rate, and beam focus), and a DOE approach was used to develop a set of 16 experiments to evaluate how these parameters influence the output. To date only one experiment was conducted for each set of parameters, so a full statistical analysis and generation of an empirical process model was not possible.

To date only one experiment was conducted for each set of parameters, so a full statistical analysis and generation of an empirical process model was not possible. However, there was enough data for the generation of plots of the marginal means and half-effects in order to evaluate the relative importance of each input parameter on the height and width of the multi-bead stacks.The results of this analysis showed that translation speed is important to both the height and width of the resulting deposit. This is partly because changing the speed at which the beam is moving changes the total volume of material that is deposited per unit length. In addition, changing the translation speed changes the amount of heat input during the process. The direct effect of this change of heat on the height and width is not fully known at this time, but metallurgical analysis shows that the samples with higher heating had a bigger melted zone.

The direct effect of this change of heat on the height and width is not fully known at this time, but metallurgical analysis shows that the samples with higher heating had a bigger melted zone. This larger melt pool is expected to cause the increase in width and decrease in height measured in the resulting multi-bead stacks.The current and voltage were also found to be important parameters to both the stackheight and width, though they were second in importance to translation speed. Increasing either the voltage or the current was found to cause an increase in the width and decrease in the height of the stack. An increase of either parameter enlarges the size of the melt pool, thereby spreading out the deposited material over a larger area and decreasing the height.114The wire feed rate was found to strongly influence the stack height but only produced a minimal change in stack width.

Increasing the wire feed rate produced an increase in the stack height, which is consistent with the larger amount of material being introduced into the melt pool. However, further analysis is required to understand the interaction between wire feed rate and stack width.The least important parameter was found to be beam focus for both stack height andwidth. Varying beam focus produced no measurable change in the stack width. In fact, metallurgical analysis confirmed that changing the beam focus produced no measurable change in the size of the melted or heat affected zones. Changing beam focus was found to produce a small change in stack height, but it had the smallest relative importance of the five parameters examined.References[1] K.M.B. Taminger and R.A. Hafley, “Electron Beam Freeform Fabrication: A Rapid Metal Deposition Process,” Proceedings of the 3rd Annual Automotive Composites Conference, Troy, MI, September 9-10 (2003)[2] W.H. Hofmeister, M.L.Griffith, M.T.

Hofmeister, M.L.Griffith, M.T. Ensz, and J.E. Smugeresky, J. E., “Solidification in Direct Metal Deposition by LENS Processing,” Journal of Metals, Vol. 53, no. 9, pp. 30-34 (2001)[3] J. Beuth, N. Klingbeil, N., “The Role of Process Variables in Laser-Based Direct Metal Solid Freeform Fabrication,” Journal of Metals, Vol. 53, no. 9, pp. 30-34 (2001)[4] S. R. Scmidt and R. G. Launsby, Understanding Industrial Designed Experiments, 2nd Edition, CQG Ltd Printing, 1989.[5] D. C. Montgomery, Design and Analysis of Experiments, 5th Edition, John Wiley and Sons Inc., 2001.[6] K.M.B. Taminger and R.A. Hafley, “Characterization of 2219 Aluminum Produced by Electron Beam Freeform Fabrication,” Proceedings of 13th SFF Symposium, 482-489 (2002)AcknowledgementsThe authors would like to express their gratitude to Jim Baughman of Lockheed Martinfor performing optical microscopy and to Erik Osborne of George Washington University for overseeing the production of the EBF3 panels.115.

SIGN OF EXPERIMENTS APPROACH FOR STATISTICAL CLASSIFICATION OF STEREOLITHOGRAPHY MANUFACTURING BUILD PARAMETERS: EFFECTS OF BUILD ORIENTATION ON MECHANICAL PROPERTIES FOR ASTM D-638 TYPE I TENSILE TEST SPECIMENS OF DSM SOMOS® 11120 RESINRolando Quintana†, Karina Puebla*, Ryan Wicker* *University of Texas at El Paso, W.M. Keck Center for 3D Innovation, El Paso, Texas 79968 †University of Texas at San Antonio, Department of Management Science and Statistics, San Antonio, Texas 78249-1644Reviewed, accepted August 21, 2007AbstractA statistical design of experiments (DOE) approach was used to determine if specific build orientation parameters impacted mechanical strength of fabricated parts. A single platform (10- inch by 10 inch cross-section) on the 3D Systems Viper si2 machine was designed to hold 18, ASTM D-638 Type I samples built in six different orientations (called Location) with three samples built for each location.

The DOE tested four factors: Location, Position, Axis, and Layout. Each sample within a Location was labeled as Positions 1, 2, or 3 depending on the distance from the center of the platform with Position 1 being the closest to the center. Samples were fabricated parallel with the x-axis, y-axis, or 45o to both axes (called Axis 1, 2, and 3, respectively) and were fabricated either flat or on an edge relative to the x-y plane (called Layout 1 and 2, respectively). The results from the statistical analyses showed that Axis, Location, and Position had no significant effect on UTS or E. However, Layout (or whether a sample was built flat or on an edge) was shown to have a statistically significant effect on UTS and E (at a 95% level of confidence). This result was not expected since a comparison of the average UTS for each Layout showed only a 1.2% difference (6966 psi versus 7050 psi for samples built flat and on an edge, respectively).

This result was not expected since a comparison of the average UTS for each Layout showed only a 1.2% difference (6966 psi versus 7050 psi for samples built flat and on an edge, respectively). Because of the small differences in means for UTS, the statistical differences between Layout most likely would not have been identified without performing the DOE. Furthermore, Layout was the only factor that tested different orientations of build layers (or layer-to-layer interfaces) with respect to the sample part, and thus, it appears that the orientation of the build layer with respect to the fabricated part has a significant effect on the resulting mechanical properties. This study represents one of many to follow that is using statistical analyses to identify and classify important fabrication parameters on mechanical properties for layer manufactured parts.

This study represents one of many to follow that is using statistical analyses to identify and classify important fabrication parameters on mechanical properties for layer manufactured parts. Although stereolithography is the focus of this work, the techniques developed here can be applied to any layered manufacturing technology.Keywords: rapid prototyping; stereolithography; design of experiments; tensile testing; WaterShed 111201. IntroductionAs the rapid prototyping (RP) or layered manufacturing (LM) industry focuses on rapid manufacturing (RM) of end-use products, the machines that were originally designed to build prototypes may now be required to build functional end-use parts.

In order to successfully accomplish this transition, the materials available for use in RP must provide the performance required for RM and the specific RP technology used for RM must be capable of providing repeatable and reproducible parts (with repeatable and reproducible performance dimensionally,mechanically, thermally, electrically, optically, etc.). Of the many additive manufacturing technologies available, stereolithography (SL) was the first technology commercialized in the 1980s and remains a very popular RP technology for building highly accurate parts with superior surface finish. The SL resin industry has also been making significant progress on improving the physical properties of photocrosslinkable resins, since these materials have traditionally performed poorly primarily in the area of impact strength.

In addition to significant investment by the industry in new better-performing resins, many research groups have been exploring ways to improve performance of SL resins. Our group, for example, has been investigating dispersing multi-walled carbon nanotubes in SL resins to improve performance (Sandoval et al., 2005 and Sandoval et al., 2006). Other groups have used electroplating to improve performance of SL- manufactured parts (Saleh et al., 2004). Assuming current and/or future SL resins will provide the performance characteristics required for a particular RM application, the question reverts back to determining the variability introduced by the manufacturing process on performance.When building RP parts, variability can be introduced in multiple ways. For example, research has shown varying mechanical performance as a result of different layer thicknesses (Chockanlingam et al., 2006).

For example, research has shown varying mechanical performance as a result of different layer thicknesses (Chockanlingam et al., 2006). If layer thickness introduces property anisostropy, the many factors associated with build orientation need to be explored. Issues associated with build orientation were explored by Hague et al. (2004) and Dulieu-Barton and Fulton (2000). In these studies, parts were built in a number of orientations and subjected to tensile testing. Both groups found differences in the tensile strength measurements for different orientations (less than 5% for Hague et al., 2004; and as much as 13% for the orientations tested by Dulieu-Barton and Fulton, 2000). Hague et al. (2004), for example, concluded that the 5% variation showed that SL produced what could be considered essentially isotropic parts for the orientations tested by them.

Hague et al. (2004), for example, concluded that the 5% variation showed that SL produced what could be considered essentially isotropic parts for the orientations tested by them. We believe, however, it is important to classify the populations using statistical analyses, and as a result, this paper explores developing a statistical design of experiments approach for classifying differences in mechanical properties resulting from different build orientations. This study investigates a rather limited parameter space, but the techniques can be applied to statistically identify and classify differences in RP-manufactured parts with respect to virtually any parameter, and our group is embarking on a considerable effort to apply these techniques to classify a number of effects introduced by varying build orientations and parameters in SL.

Furthermore, although SL is the focus of this work, the techniques developed and demonstrated here can be applied to any layered manufacturing technology.2. Experimental MethodsMaterialsA commercially available epoxy-based resin, WaterShed™ 11120 (DSM Somos®, Elgin, IL)was used to perform this investigation. WaterShed™ 11120 is a widely used multi-purpose resin, characterized by a low viscosity (~ 260 cps at 30 º C) with a density of ~1.12g/cm3 at 25 º C (Product Data Sheet, DSM Somos®). The epoxy-based resin is used for solid-state laser systems (354.7 nm laser wavelength) with a recommended critical exposure, Ec, of ~11.5 mJ/cm2 and penetration depth, Dp, of ~6.5 mJ/cm2. Using the recommended settings, a 3D System Viper si2 SL machine manufactured all test specimens presented here.MethodsASTM D-638 Type I tensile test specimens (ASTM, 2005) as shown in Figure 1 and Table 1were manufactured for this study.

The thickness of 0.13-inches applies to molded plastics andwas used here. A separate study showed that the Viper si2 was a capable process for producing repeatable and reproducible parts using the ASTM tolerances shown in Table 1 (these results are provided separately in Quiñones et al., 2006).Figure 1. ASTM D-638 Type I tensile test specimen.Table 1. ASTM D-638 Type I tensile test specimen dimensions and tolerances.Dimensions (in)Type ITolerancesW, (Wc) - Width of narrow section L - Length of narrow section WO - Width overall LO - Length overall G - Gage length D - Distance between grips R - Radius of fillet T - Thickness0.5 2.25 0.75 6.5 2 4.5 3 0.13± 0.02 ± 0.02 + 0.25 no max ± 0.01 ± 0.2 ± 0.04 ± 0.02The layout of build orientations tested in the DOE is provided in Figure 2 and the levels for each factor are shown in Table 2.

A single platform (10-inch by 10 inch cross-section) on the 3D Systems Viper si2 machine was designed to hold 18, ASTM D-638 Type I samples and the experimental design tested four factors, called Location, Position, Axis, and Layout. The 18 total samples were divided into six platform Locations with three samples per each Location. Each sample within a Location was labeled as Positions 1, 2, or 3 depending on the distance from the center of the platform with Position 1 being the closest to the center. Samples were fabricated parallel with the x-axis, y-axis, or 45o to both axes (called Axis 1, 2, and 3, respectively) and were fabricated either flat or on an edge relative to the x-y plane (called Layout 1 and 2, respectively). This layout enabled a single platform to produce 18 samples with 3 replicates for each factor. A single build was used in the DOE to remove any batch to batch variations and aging effects.

This layout enabled a single platform to produce 18 samples with 3 replicates for each factor. A single build was used in the DOE to remove any batch to batch variations and aging effects. As will be described in the results, there were sufficient samples in this approach to make statistical conclusions.Table 2. Specific factors with levels tested in DOE. AxisLocation PositionLayout Flat Flat Flat On Edge On Edge On Edge1 2 3 4 5 61 2 3 1 2 3x y xy x y xy1 2 3 1 2 31 1 1 2 2 2The test specimens were drawn using SolidWorks as shown in Figure 2, and then converted to stl file format. The stl file was processed with 3D Lightyear, sliced into standard 0.004-inch thick layers, and built using the Viper si2 (laser power of ~45 mW) with the resin manufacturer’s recommended build parameters. The build time for the 18 test specimens was 9 hours.

The build time for the 18 test specimens was 9 hours. Once finished, the samples were removed from the platform, cleaned with isopropyl alcohol, and post- cured inside an ultraviolet oven for 30 minutes on each side. After post-curing, any remaining support material was removed by hand. Prior to pulling the test specimens on an Instron 5866 tensile testing machine (using a 10 kN static load cell with a cross-head test speed of 0.2 in/min), each sample was measured to ensure the samples were manufactured within the ASTM specifications.Figure 2. Layout of the Viper si2 build platform illustrating 18 samples in 6 different orientations with 3 replicates for each orientation.3.

Layout of the Viper si2 build platform illustrating 18 samples in 6 different orientations with 3 replicates for each orientation.3. Tensile Test ResultsMechanical TestingThe 18 samples shown in Figure 2 were subjected to tensile testing (in random order) using an Instron 5866 tensile testing machine and Ultimate Tensile Stress (UTS, in psi), Elongation at Break (EB, in %), Fracture Stress, (FS, in psi), and Young’s Modulus (E, in psi) were measured and recorded. The results are shown in Table 3 and Figure 3. The minimum standard deviations for the measurements occur for UTS and E, as expected, and thus, these measurements are used in the statistical analyses in the following sections.Table 3.

Stress vs. strain results (Data split into two figures for visual clarity).Top: The first nine random specimens tested, and Bottom: The second nine random specimens tested.4. Experimental DesignUsing the results from the tensile testing as shown in Table 3 and Figure 3, the statisticalanalyses were performed. The following section describes and provides rationale for the experimental design and presents the experimental results and concomitant statistical analysis. The key assumptions underlying the adequacy of the statistical models are then presented.The tensile strength completely randomized experiment on ASTM Type I specimensinvolved the study of the effects of three factors (Axis, Layout, and Position) on ultimate tensile strength (UTS) and Young’s modulus (E).

This preliminary investigation was focused on answering the following questions:What effects do the geometric configuration factors (i.e., Axis, Position, and Layout) have on tensile strength (ultimate tensile strength - UTS and Young’s modulus – E) for an ASTM Type I specimen?Is there a geometric configuration that would give uniformly higher tensile strength foran ASTM Type I specimen?In each complete replication of the experiment, all possible combinations of the levels of the factors (3 for Axis, 2 for Layout, and 3 for Position) were investigated. All 18 parts were made from the same build (batch). Therefore, a three-factor factorial model was first developed, using Axis, Position and Layout as factors with UTS and E as dependent variables.

All 18 parts were made from the same build (batch). Therefore, a three-factor factorial model was first developed, using Axis, Position and Layout as factors with UTS and E as dependent variables. The three-factor analysis of variance model is provided in Equation 1.y ijkl) µ τ β γ τβ + j(= + ++ikij+( ) τγik+() βγjk+( ) τβγijk+εijkli =⎧ ⎪ =⎪ j ⎨ k = ⎪ ⎪ l = ⎩1, 2,31, 2,31, 21where τ represents Position factor with 3 levels, β the Axis factor with 3 levels, and γ the Layout factor with 2 levels constituting 3x3x2 = 18 complete cases made in random order. Note that ε the random error component, has a single level corresponding to a single replicate per cell due the destructive nature of the tensile strength testing.Location was studied independently of other factor since it is a linear combination due to physical restrictions.

For example, Locations 1, 2, and 3 only correspond to Layout 1, while Locations 4, 5, and 6 correspond to Layout 2. Without loss of continuity, future investigation may involve a nested or hierarchical design, with the corresponding levels of the Location factor under the levels of factor Layout under a two-staged nested design. In the meantime, the one- factor ANOVA model is provided in Equation 2.y ij= + +µ τ ε i iji =⎧ ⎨ ⎩1, 2,3, 4,5, 6j=1, 2,3In each of three (j) complete replications of the experiment, all possible combinations of the6 Location levels (i) were investigated. Again all 18 parts were made from the same build (batch).4.1 Experimental ResultsTables 4 and 5 present the experimental results for Position, Axis and Layout by UTS and E,respectively.12Table 4.

Results for UTS by orientation.LayoutPosition 1 2 31 7071.3 6903.7 7014.91 Axis 2 6991.8 6972.7 6924.83 6954.2 6841.5 7021.51 7022.8 7028.9 6980.22 Axis 2 7080.3 7076.6 7005.73 7029.5 7182.4 7040.3Table 5. Results for E by orientation.LayoutPosition 1 2 31 121722.1 120890.3 121222.71 Axis 2 123736.8 121897.1 121214.83 121710.7 119533.5 121007.81 125495.8 123678.5 121844.42 Axis 2 120311.0 122100.2 123914.13 124550.6 124916.9 123656.7Tables 6 and 7 show the mean UTS and E, respectively, for each level of the factors. It alsoshows the standard error of each mean, which is a measure of its sampling variability. The rightmost two columns show 95.0% confidence intervals for each of the means.Table 6. Least squares means for UTS with 95.0 percent confidence intervals. Stnd.

The rightmost two columns show 95.0% confidence intervals for each of the means.Table 6. Least squares means for UTS with 95.0 percent confidence intervals. Stnd. Lower Upper Level Count Mean Error Limit Limit GRAND MEAN 18 7007.94 Position 1 6 7024.83 29.5885 6960.37 7089.3 2 6 7001.17 29.5885 6936.7 7065.63 3 6 6997.83 29.5885 6933.37 7062.3 Layout 1 9 6966.33 24.1589 6913.7 7018.97 2 9 7049.56 24.1589 6996.92 7102.19 Axis 1 6 7003.67 29.5885 6939.2 7068.13 2 6 7008.83 29.5885 6944.37 7073.3 3 6 7011.33 29.5885 6946.87 7075.8Table 7. Least squares means for E with 95.0 percent confidence intervals. Stnd.

Least squares means for E with 95.0 percent confidence intervals. Stnd. Lower Upper Level Count Mean Error Limit Limit GRAND MEAN 18 122411.0 Position 1 6 122921.0 636.302 121535.0 124308.0 2 6 122169.0 636.302 120783.0 123556.0 3 6 122144.0 636.302 120757.0 123530.0 Layout 1 9 121437.0 519.539 120305.0 122569.0 2 9 123385.0 519.539 122253.0 124517.0 Axis 1 6 122476.0 636.302 121089.0 123862.0 2 6 122196.0 636.302 120809.0 123582.0 3 6 122563.0 636.302 121177.0 123949.04.2 Multifactor Analysis – Position, Axis, LayoutA multifactor ANOVA for the dependent variables UTS and E, respectively, by levels of the Position, Axis and Layout factors is first presented.

Various tests and graphs to determine which factors have a statistically significant effect on the dependent variables were performed, as well as tests for significant interactions amongst the factors.The multifactor ANOVA for UTS and E, summarized in Tables 8 and 9, respectively, decomposes the variability of the dependent variable (UTS and E) into contributions due to the various factors. All F-ratios are based on the residual mean square error. Further, Type III sums of squares was utilized since it allows for the contribution of each factor to be measured after having removed the effects of all other factors.Table 8. Analysis of variance for UTS - Type III sums of squares. Source Sum of Squares Df Mean Square F-Ratio P-Value MAIN EFFECTS A:Position 2600.44 2 1300.22 0.25 0.7846 B:Layout 31166.7 1 31166.7 5.93 0.0314 C:Axis 183.444 2 91.7222 0.02 0.9827RESIDUAL 63034.3 12 5252.86 TOTAL (CORRECTED) 96984.9 17Table 9. Analysis of variance for E - Type III sums of squares.

Analysis of variance for E - Type III sums of squares. Source Sum of Squares Df Mean Square F-Ratio P-Value MAIN EFFECTS A:Position 2.34239E6 2 1.1712E6 0.48 0.6289 B:Layout 1.70742E7 1 1.70742E7 7.03 0.0211 C:Axis 441793.0 2 220897.0 0.09 0.9137RESIDUAL 2.91514E7 12 2.42928E6 TOTAL (CORRECTED) 4.90098E7 17Tables 10 - 15 apply a multiple comparison procedure to determine which means aresignificantly different from which others for UTS and E by Position, Axis and Layout, respectively. The bottom half of the output shows the estimated difference between each pair of means. Within each column, the levels containing X's form a group of means within which there are no statistically significant differences. The method used to discriminate among the means is Fisher's least significant difference (LSD) procedure. With this method, there is a 5.0% risk of calling each pair of means significantly different when the actual difference equals 0.Table 10.

With this method, there is a 5.0% risk of calling each pair of means significantly different when the actual difference equals 0.Table 10. Multiple range tests for UTS by Position. Method: 95.0 percent LSD Position Count LS Mean LS Sigma Homogeneous Groups 3 6 6997.83 29.5885 X 2 6 7001.17 29.5885 X 1 6 7024.83 29.5885 X Contrast Difference +/- Limits 1 - 2 23.6667 91.1713 1 - 3 27.0 91.1713 2 - 3 3.33333 91.1713Table 11. Multiple range tests for E by Position.

Multiple range tests for E by Position. Method: 95.0 percent LSD Position Count LS Mean LS Sigma Homogeneous Groups 3 6 122144.0 636.302 X 2 6 122169.0 636.302 X 1 6 122921.0 636.302 X Contrast Difference +/- Limits 1 - 2 752.0 1960.65 1 - 3 777.833 1960.65 2 - 3 25.8333 1960.65There are no statistically significant differences between the mean values of UTS and E from one level of Position to another at a level of confidence of 95%, as depicted in Figures 4a and 4b.Means and 95.0 Percent LSD IntervalsMeans and 95.0 Percent LSD Intervals7080(X 1000) 1247050123.57020123S T U6990E122.569601226930121.5690012112 Position312 Position3Figure 4a. Means and 95% LSD, UTS by PositionFigure 4b. Means and 95% LSD, E by Positionintervals.intervals.Table 12. Multiple range tests for UTS by Axis.

Multiple range tests for E by Axis. Method: 95.0 percent LSD Axis Count LS Mean LS Sigma Homogeneous Groups 2 6 122196.0 636.302 X 1 6 122476.0 636.302 X 3 6 122563.0 636.302 X Contrast Difference +/- Limits 1 - 2 279.833 1960.65 1 - 3 -87.5 1960.65 2 - 3 -367.333 1960.65There are no statistically significant differences between the mean values of UTS and E fromone level of Axis to another at a level of confidence of 95%, as depicted in Figures 5a and 5b.Means and 95.0 Percent LSD IntervalsMeans and 95.0 Percent LSD Intervals7070(X 1000) 1247050123.57030123S T U7010E122.569901226970121.5695012112 Axis312 Axis3Figure 5a. Means and 95% LSD, UTS by AxisFigure 5b. Means and 95% LSD, E by Axisintervals.intervals.Table 14. Multiple range tests for UTS by Layout.

Means and 95% LSD, UTS by AxisFigure 5b. Means and 95% LSD, E by Axisintervals.intervals.Table 14. Multiple range tests for UTS by Layout. Method: 95.0 percent LSD Layout Count LS Mean LS Sigma Homogeneous Groups 1 9 6966.33 24.1589 X 2 9 7049.56 24.1589 X Contrast Difference +/- Limits 1 - 2 *-83.2222 74.4411 * denotes a statistically significant difference.Table 15. Multiple range tests for E by Layout.

Multiple range tests for E by Layout. Method: 95.0 percent LSD Layout Count LS Mean LS Sigma Homogeneous Groups 1 9 121437.0 519.539 X 2 9 123385.0 519.539 X Contrast Difference +/- Limits 1 - 2 *-1947.89 1600.86 * denotes a statistically significant difference.An asterisk has been placed next to 1 pair in both preceding tables, indicating that there arestatistically significant differences between the mean values of UTS and E from one level of Layout to another at a level of confidence of 95%, as depicted in Figures 6a and 6b.Means and 95.0 Percent LSD IntervalsMeans and 95.0 Percent LSD Intervals7100(X 1000) 1257060124S T U70206980E1231226940121690012120Layout1Layout2Figure 6a. Means and 95% LSD, UTS by LayoutFigure 6b. Means and 95% LSD, E by Layoutintervals.intervals.Tables 16 and 17 summarize the significance of factor levels versus UTS and E, respectively,at a level of confidence of 95%.

Means and 95% LSD, E by Layoutintervals.intervals.Tables 16 and 17 summarize the significance of factor levels versus UTS and E, respectively,at a level of confidence of 95%. Layout (P-value ~ 0.0314 for UTS and ~0.0211 for E) was the only orientation that had a statistically significant effect on the respective dependent variables at the 95.0% level of confidence. The physical significance of this result is that the individual build layers (x-y plane) is the same for Axis, Location, and Position. Layout is the only factor that has individual layers in a different orientation with respect to the test samples. It is also important to note that the vertical orientation (where the L for the Type I sample is built vertically) was not tested and based on these results, we would expect a difference for this orientation as well. This experiment is left for future work.Table 16. Significance summary of UTS by factor at the 95% level of confidence.

This experiment is left for future work.Table 16. Significance summary of UTS by factor at the 95% level of confidence. FactorAxisLayout Position 1 2 31 2 3 1 N N N2Axis1 2 3 1 2 Position 1 2 3LayoutYN N NTable 17. Significance summary of E by factor at the 95% level of confidence. FactorAxisLayout Position 1 2 31 2 3 1 N N N2Axis1 2 3 1 2 Position 1 2 3LayoutYN N N4.3 One Factor Analysis – LocationVarious tests and graphs to determine which Location levels have a statistically significanteffect on the dependent variables UTS and E were performed, as well as tests for significant interactions amongst the levels of Location. As shown in the results for UTS and E in Tables 18 and 19, 3 replicates for each of the 6 levels of Location for a total of 18 observations per dependent variable were run.Table 18.

As shown in the results for UTS and E in Tables 18 and 19, 3 replicates for each of the 6 levels of Location for a total of 18 observations per dependent variable were run.Table 18. Results for UTS by Location.Location 1 2 3 4 5 61 7071.3 6991.8 6954.2 7022.8 7080.3 7029.5Observations 2 6903.7 6972.7 6841.5 7028.9 7076.6 7182.43 7014.9 6924.8 7021.5 6980.2 7005.7 7040.3Table 19. Results for E by Location. Observations 2 120890.3 121897.1 119533.5 123678.5 122100.2 124916.9Location 1 2 3 4 5 61 121722.1 123736.8 121710.7 125495.8 120311.0 124550.63 121222.7 121214.8 121007.8 121844.4 123914.1 123656.7Tables 20 and 21 show the mean UTS and E, respectively, for each level of the factors. It also shows the standard error of each mean, which is a measure of its sampling variability. The rightmost two columns show 95.0% confidence intervals for each of the means.Table 20. Summary statistics for UTS by Location. Location Count Average Variance Stnd. dev.

Summary statistics for E by Location. Location Count Average Variance Stnd. dev. Minimum 1 3 121278.0 175352.0 418.751 120890.0 2 3 122283.0 1.70187E6 1304.56 121215.0 3 3 120751.0 1.23437E6 1111.02 119534.0 4 3 123673.0 3.3343E6 1826.01 121844.0 5 3 122108.0 3.24545E6 1801.51 120311.0 6 3 124375.0 420132.0 648.176 123657.0 Total 18 122411.0 2.88293E6 1697.92 119534.0 Location Maximum Range Stnd. skewness Stnd. kurtosis 1 121722.0 832.0 0.413122 2 123737.0 2522.0 0.859076 3 121711.0 2177.0 -0.696667 4 125496.0 3652.0 -0.00929374 5 123914.0 3603.0 0.0147187 6 124917.0 1260.0 -0.800305 Total 125496.0 5962.0 0.486497 -0.805102There is more than a 3 to 1 difference between the smallest standard deviation and the largest. This may cause problems since the analysis of variance assumes that the standard deviations at all levels are equal, and the error term is normal and identically distributed (NID).

This may cause problems since the analysis of variance assumes that the standard deviations at all levels are equal, and the error term is normal and identically distributed (NID). A variance check presented in the Model Adequacy section indicates that there is not a statistically significant difference amongst the standard deviations at the 95.0% confidence level. Nevertheless, we choose to err on the conservative side in this case where the normality assumption may not be valid and utilize an alternative to the F-test procedure presented in Equation 2. The non-parametric Kruskal-Wallis test (Kruskal and Wallis, 1952) is thus utilized to compare medians rather than means.The yij observations are first ranked in ascending order, and each observation is then replacedby its rank (Rij), with the smallest observation assigned a rank = 1. In case of ties the average rank is assigned to each tied observation.

In case of ties the average rank is assigned to each tied observation. If Ri = ∑ (the ranks in the ith treatment or Location), then the test statistic is shown in Equation 3, where ni = 3 (the number of observations in the ith Location) and N = 18 (total number of observations).H⎡ 1 = ∑ ⎢ 2 S ⎣6i1 =2 R i ni−( NN + 4) 12⎤ ⎥ ⎦S2, the variance of the ranks, is given in Equation 4, and barring ties the test statistic H is presented in Equation 5.S2=N1 −1⎡ ⎢ ⎣63∑∑1 =j1 =2 R i−( NN + 4) 12⎤ ⎥ ⎦H=⎡ ⎢ ⎣12 ( NN+) 16∑i1 =2 R i ni⎤ −⎥ ⎦( 3N+)1Tables 22 and 23 present the results of the Kruskal-Wallis test for UTS and E by Location, respectively, where the null hypothesis is that the medians of the dependent variables within each of the 6 levels of Location are the same.

Box and whisker plot, E by Locationintervals.intervals.4.4 Model AdequacyIn all models generated no indications were present to assume that the fundamentalconditions for the models were violated, based on the following analyses:The normal probability plots do not reveal violations of the normal and identically distributed (NID) error term.The residuals are structure less; that is, the plot of residuals versus fitted values of the dependent variables did not reveal any obvious patterns.Standardized residuals were approximately normal with mean = 0 and unit variance – no residual was greater than 2 standard deviations, and thus outliers did not seriously distort the ANOVA.The residual plots showed no significant tendencies, and thus no correlation between residuals.No 2-factor interactions in the models were present.The four statistics displayed in Tables 24-26 test the null hypothesis that the standarddeviations of the dependent variable (UTS and E) within each of the levels of each of the factors, respectively, are the same (Montgomery, 1997).

Of particular interest are the P-values. Since the smallest of the P-values is greater than or equal to 0.05, there is not a statistically significant difference amongst the standard deviations at the 95.0% confidence level. As a result, the important assumption that the variance of the observations did not change significantly as the magnitude of the observation changed is upheld. It should be noted that the measuring instruments were calibrated before each test.Table 24. Variance check – UTS by Factor.Factor(cid:198)PositionLayoutAxisTestResult Cochran's C 0.492 Bartlett's 1.044 Hartley's 3.012 Levene's 1.154P-Value 0.71 0.600.33Result 0.580 1.013 1.384 0.489P-Value 0.65 0.65 N/A 0.49Result 0.643 1.220 3.853 0.632P-Value 0.11 0.25 N/A 0.54Table 25.

A single platform (10-inch by 10 inch cross-section) on the 3D Systems Viper si2 machine was designed to hold 18, ASTM D-638 Type I samples in six different orientations (called Location) with three replicates for each Location. Four factors were tested in the DOE: Location, Position, Axis, and Layout. Each sample within a Location was labeled as Positions 1, 2, or 3 depending on the distance from the center of the platform. Samples were fabricated parallel with the x-axis, y- axis, or 45o to both axes (called Axis 1, 2, and 3, respectively) and were fabricated either flat or on an edge relative to the x-y plane (called Layout 1 and 2, respectively). The results from the statistical analyses showed that Axis, Location, and Position had no significant effect on UTS or E. However, Layout (or whether a sample was built flat or on an edge) was shown to have a statistically significant effect on UTS and E (at a 95% level of confidence).

However, Layout (or whether a sample was built flat or on an edge) was shown to have a statistically significant effect on UTS and E (at a 95% level of confidence). Physically, this result can be explained by the orientation of individual layers (or layer-to-layer interfaces) with respect to the part, since Axis, Location, and Position all had identical sample layer interfaces. That is, Layout was the only factor that tested different orientations of layer interfaces with respect to the test samples (samples built flat versus those built on an edge produced orthogonal layer interfaces with respect to the part). Without performing the DOE, the statistical differences between Layout most likely would not have been identified (since a comparison of the average UTS for each Layout showed only a 1.2% difference).

Should one conclude that building a part flat or on an edge produces parts with essentially isotropic properties (based on the 1.2% difference in the means of UTS for the samples), this conclusion would be statistically incorrect. As a result of this work, it appears that the orientation of the individually fabricated layers (or the layer interfaces) with respect to the part produces samples with statistically different mechanical properties (UTS and E in this case). This study did not include a test of samples built vertically (where the major axis of the part is built vertically or in the z direction of the machine). Based on the present work, we would expect vertically built samples to have UTS and E properties that are statistically different from samples built flat and on an edge. Testing this hypothesis statistically will be left for future work.AcknowledgmentsThe facilities within the W.M.

Testing this hypothesis statistically will be left for future work.AcknowledgmentsThe facilities within the W.M. Keck Center for 3D Innovation (Keck Center) used here contain equipment purchased through Grant Number 11804 from the W.M. Keck Foundation, a faculty STARS Award from the University of Texas System, and two equipment grants from Sandia National Laboratories. This material is based in part upon work supported through the Mr. and Mrs. MacIntosh Murchison Chair I in Engineering and through research contract 504004 from Sandia National Laboratories in the Laboratory Directed Research and Development (LDRD) program. The authors are grateful for the assistance of Luis Ochoa and Jesus Flores within the Keck Center for their assistance during performance of various aspects of the research.

The authors are grateful for the assistance of Luis Ochoa and Jesus Flores within the Keck Center for their assistance during performance of various aspects of the research. Sandia National Laboratories is a multi-program laboratory operated by Sandia Corporation, a Lockheed Martin Company, for the United States Department of Energy’s National Nuclear Security Administration under contract DE-AC04-94AL85000.References1. American Society for Testing and Materials, Standard 638 “Standard Test Method for Tensile Properties of Plastics,” Annual Book of ASTM Standards, West Conshohocken, PA, 2005.2. Chockalingam, K., Jawahar, N., “Influence of Layer Thickness on Mechanical Properties in Stereolithography,” Rapid Prototyping Journal, 2006, Volume 12 Number 2, pp. 106-113.3.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDESIGN OF FIBER-REINFORCED CELLULAR STRUCTURES WITH TENSEGRITY BEHAVIOR MANUFACTURED USING 3D PRINTED SAND MOLDSNikhil Jorapur1, 2, Robert West2, Christopher B. Williams1, 2, Alan Druschitz31 Design, Research and Education for Additive Manufacturing Systems Lab 3 Virginia Tech Foundry Institute for Research and Education 2 Department of Mechanical Engineering 3 Department of Material Science & Engineering Virginia Tech, Blacksburg, VA 24060AbstractTensegrity structures provide a high stiffness to mass ratio since all the comprising elements are either in compression or tension.

However, they have limited applications since fabrication of such structures is challenging due to their complexity and mainly requires manual assembly of components. The authors look to Additive Manufacturing (AM) as a means to introduce tensegrity behavior in cellular structures to enhance structural performance. Specifically, octet cellular structures are created by casting aluminum into 3D-printed sand molds embedded with continuous wires. In this paper, the authors describe design and analysis of octet cellular structures that feature high strength fibers held in continuous tension. Finite element analysis of 4- point bending test is used to evaluate the effectiveness of embedded fibers. Also, the presence of tensegrity behavior was evaluated using this analysis and testing. The structure with tensegrity behavior was found to be 30 % stronger.

Also, the presence of tensegrity behavior was evaluated using this analysis and testing. The structure with tensegrity behavior was found to be 30 % stronger. The simulation and experimental results were shown to match within 6 % error in the elastic region.1Introduction1.1 Cellular MaterialsCellular materials are a set of lightweight structures, which provide sufficient stiffness and strength with a lower mass density for a particular loading [1] [2]. These structures are of great interest due to their efficient structural performance and have been developed for a wide range of applications in automotive, bio-medical, aerospace, sporting, construction and energy-absorption systems [3]. Cellular structures are characterized by a network of cells comprising inter-connected struts, which help in reducing material present in the structure [2]. They are differentiated by the nature of the voids present in the structure.

They are differentiated by the nature of the voids present in the structure. Stochastic materials are composed of cells of random shape, morphology and distribution [4]. Ordered/periodic cellular materials are composed of an ordered repetition of standard cell topology [5]. Cellular materials featuring designed mesostructures are differentiated by their selective placement of material to achieve multiple design objectives (i.e., the topology is repeatable and controlled but not necessarily periodically repeating) [5].Recent efforts have been directed to improve performances of cellular structures. Using stiffer, higher strength materials, compressive stiffness and strength can be increased. Through fabrication of lattices at millimeter scale it is shown that the cellular structure can withstand higher axial stresses [6], [7], [8].

Through fabrication of lattices at millimeter scale it is shown that the cellular structure can withstand higher axial stresses [6], [7], [8]. This is simply because materials gain strength at lower scales, which implies high elastic modulus in structures can make them significantly stronger.2043Additive Manufacturing (AM) has enabled progress in improving processes, design and materials for fabricating cellular materials with designed mesostructures. Electron beam melting (EBM) process was used to manufacture Ti-6Al-4V auxetic structures with negative Poisson’s ratio, which were shown to be superior compared to regular foam structures [9]. Projection microstereolithography was used to produce octet microlattices exhibiting ultrastiff properties without being affected by density [10].

Projection microstereolithography was used to produce octet microlattices exhibiting ultrastiff properties without being affected by density [10]. In this paper, the authors aim to increase the bending stiffness of octet-cellular structures through incorporation of tensegrity behavior realized by reinforcement of continuous high strength fibers. In this context, AM is used to solve the manufacturing challenges to realize tensegrity behavior.Fiber-reinforced cellular composite structures have been proposed as a means of solving structural limitations of cellular materials and bridging the space between current materials and unachievable low-density region in Ashby’s chart [11]. Micro-buckling failure in struts of pyramidal truss structures was improved through carbon-fiber reinforced polymer laminates [12]. Unidirectional carbon fiber/epoxy prepreg laps were hot-pressed to form pyramidal structures with superior compressive strength. [13].

Unidirectional carbon fiber/epoxy prepreg laps were hot-pressed to form pyramidal structures with superior compressive strength. [13]. Sandwiched composite foldcores with carbon-reinforcement were produced to enhance the brittleness of the structures enabling better energy-absorption [14]. Hollow composite trusses composed of carbon-fibers were arranged to form pyramidal truss structures with high buckling strength using thermal expansion molding [15]. Composite cellular materials were processed by reversibly assembling and integrating looped fiber composite beams in a millimeter scale, which extended stretch-dominated lattice structures to ultra-light regime [16].

Silicon-carbide filaments were reinforced inside the titanium struts to fabricate millimeter scale lattice structures outperforming other cellular materials with density of less than 1 g cm-3 [17], [18].1.2 Context: Tensegrity StructuresTensegrity arrangement is designed to involve only tensile and compression stresses regardless of loading style [19]. The word tensegrity is a reduction of tensile integrity, coined by R. B. Fuller through his patent in 1962 [20]. The structure is made of spatial assemblies of rigid compressive bars and deformable pre-stressed strings [21]. Their non-linear mechanical behavior along with minimal mass formations [22] for given loads make them suitable for lightweight deployable applications. Tensegrity structures can also absorb and distribute energy under different load conditions [23], [24]. The strings help in distribution of compressive stress of rods through their pre-tension and high strength.

The strings help in distribution of compressive stress of rods through their pre-tension and high strength. A robotic leg was created using the tensegrity arrangement of bone-tendon such that pre-tensioned wires were included to absorb energy and enable high speed running [25]. However, their applications are limited due to the manufacturing challenges caused by complexity and are often fabricated by manual assembly. In order to address this gap, the authors present the use of a process chain of Additive Manufacturing (AM) and metal- casting.The effort towards fiber-reinforcement of cellular materials can be approached through a tensegrity perspective. In this paper, authors aim to embed continuous long fibers such that tensegrity behavior is introduced in cellular structure, consequently improving the structural performance.

In this paper, authors aim to embed continuous long fibers such that tensegrity behavior is introduced in cellular structure, consequently improving the structural performance. The study reported in this paper concentrates on developing performance of octet cellular structures for bending applications.20441.3 Context: Octet Cell StructureMost of the instances where bending loads are subjects, cellular structures are significantly used [6], [7], [26], [27]. Cellular structures with stretch dominated cores have been produced to predominantly experience axial stresses when loaded in bending [6], [18]. Also, their performance is considered to be superior to panels, whose struts deform by bending. Octet unit cell, as shown in Figure 1 is considered to be stretch dominated core where predominantly axial stresses are experienced under bending loads [1].

Octet unit cell, as shown in Figure 1 is considered to be stretch dominated core where predominantly axial stresses are experienced under bending loads [1]. These cores which experience axial stress in bending can be complemented by fiber, which can act as pre-stressed strings to introduce tensegrity behavior, making the structures more efficient. Also, it would be nearly pure tensegrity behavior, since there would be little presence of bending stress in such core structures [6].Figure 1. Octet 3D cell structureIn this paper the authors present the method of introducing the tensegrity behavior in octet cellular structures through finite element analysis. The authors describe the design of structure and fabrication process made possible through AM and traditional casting techniques. Section 3 describes the concept of locating fibers such that they are loaded in tension. Section 4 discusses the modeling of octet cellular structure with fibers following methods from Section 2.

Section 3 describes the concept of locating fibers such that they are loaded in tension. Section 4 discusses the modeling of octet cellular structure with fibers following methods from Section 2. The tensegrity behavior is also studied in this section. Section 4 describes the fabrication methods and experimental testing on the structures. Results are discussed in last section with a closure of the work reported in this paper.2 Background and Motivation2.1 Limitations of Cellular MaterialsWhile cellular materials provide superior structural performance, their applications are constrained by lower elastic moduli and indentation weakness [2] [28]. These structures are predominantly affected by concentrated compressive stresses [29]. Also, the core of the panel suffers local yield or crushing accompanied with face sheet stretching/tearing [30], [31], [32].

Also, the core of the panel suffers local yield or crushing accompanied with face sheet stretching/tearing [30], [31], [32]. All these limitations are reported to be affected by strength of the core and cell size [33].2.2 Continuous Fiber-ReinforcementNature provides best examples in exhibiting mechanical efficiency where often continuous long fibers are observed to enhance structural performance. Investigation on pomelo fruit, which can withstand a drop of 10 m without significant damage, has revealed that the peel consists of2045hierarchical cellular structures with fibrous supports that boost energy absorption, as shown in Figure 2 [34]. In an effort to create a man-made structure with similar properties, the same research group attempted to produce metal-metal composite structure to mimic pomelo’s behavior.

In an effort to create a man-made structure with similar properties, the same research group attempted to produce metal-metal composite structure to mimic pomelo’s behavior. Aluminum with two different variants of ductility, outside shell being less ductile was constructed and evaluated for energy absorption [35].Fibrous BundlesFigure 2. SEM images of pomelo peel samples [34]In this paper, the work has been inspired by the presence of continuous fibers in the aforementioned structures, where performance in terms of energy-absorption and strength has been enhanced. Using these continuous fibers authors seek to introduce tensegrity behavior and provide another answer to improve cellular structures. For this solving manufacturing challenges imposed by tensegrity structures is important. Actuated joints were created by embedding shape memory alloys using Polyjet 3D printing process [36]. Copper wire was embedded in fused deposition modeling process using ultrasonic energy [37].

Actuated joints were created by embedding shape memory alloys using Polyjet 3D printing process [36]. Copper wire was embedded in fused deposition modeling process using ultrasonic energy [37]. Figure 3 represents the basic approach followed in this paper. For embedding fiber in any AM system, the print needs to be paused to embed fiber and continue printing to attain the final part.Pause printingEmbed fiberResume printingFigure 3. Schematic representation of embedding fiber while 3D printing20463 Modeling and AnalysisThe purpose of modeling was to make design decisions such as truss dimensions, placement of fiber and validating tensegrity behavior in cellular structures.

Finite element analysis on two-dimensional truss arrangements and octet 3D beams structure was conducted in ABAQUS/Explicit [38] to produce a design of cellular structures with tensegrity behavior.In order to integrate tensegrity behavior it is important to determine the placement of fiber, such that the load distribution occurs due to the tensioning of fibers. This study was expanded and applied to octet-cellular structure. The presented analysis technique can be effectively used to include tensegrity behavior in other cellular materials.For the analysis studies, Aluminum 6061-T6 was chosen as the material for trusses and cellular structures. The material was assumed to isotropic, homogeneous and hookean for elastic deformations. The material was given an elastic modulus of 68.9 GPa [39] and Poisson’s ratio of 0.33 [39]. The structures are subjected to 4-point bending test since the structure is being designed for a bending application.

The structures are subjected to 4-point bending test since the structure is being designed for a bending application. Bending performance is evaluated in terms of stresses and displacements occurring in the elastic region.3.1 ConceptA two-dimensional arrangement of trusses was created to provide basis of the concept. Figure 5 (a) depicts the structure, which was studied to establish the concept of including high- strength fibers to achieve tensegrity mechanics in cellular structures. This study was an inexpensive method to represent cellular structures and provide guidance while embedding fibers.3.1.1 Truss DimensionsUsing the curve Euler-Johnson curve for aluminum as reference, the truss dimensions were chosen such that the struts did not fail due to buckling. Strut diameter was chosen to be 7 mm to ensure the complete filling of mold while casting and also allow enough bonding with fiber.

Strut diameter was chosen to be 7 mm to ensure the complete filling of mold while casting and also allow enough bonding with fiber. It is essential for bonding to occur so that the wires are engaged completely in tension. The wires would otherwise slip in their place and not contribute towards reduction of stress concentration. For a slenderness ratio of 24.5, the strut length was chosen to be 25 mm.3.2 Truss 2D Finite Element ModelingThe structure, as shown in Figure 5 (a), was subjected to 4-point bending test where the loads were applied parametrically and the displacement of the center point was noted for each loading condition. The truss-structure were all represented by 2D beam elements – B21, the Timoshenko beam elements meant for short columns. The loading and boundary conditions were set for 4-point bending tests. The loading points are indicated Figure 4.

The loading and boundary conditions were set for 4-point bending tests. The loading points are indicated Figure 4. One end of the structure was pinned while the other end was allowed to move only in X-direction (roller) as shown in Figure 5 (a).2047(a)RollerPinned121314151617(b)1117891062345Figure 4. 2D Truss a) loading and boundary conditions, b) deformed state after loading3.2.1 ResultsThe displacement of the points in the 2D Truss structure was studied. Nodes 1, 2, 3, 7, 8 and 9, as shown in Figure 5, were considered since the structure is symmetric and the displacements would be the same for the corresponding mirrored nodes. Also the nodes on the top will have the negative displacement with the same magnitude as the corresponding bottom nodes.Figure 5. Displacements (orthogonal to loading direction) at selected nodesPoints where displacement occurred orthogonally to the loading direction were noted.

Displacements (orthogonal to loading direction) at selected nodesPoints where displacement occurred orthogonally to the loading direction were noted. Nodes 1, 2 and 3 show high displacements, which means nodes its true for nodes 4, 5 and 6 as well. These points are candidates for continuous fiber reinforcement. If a fiber were to be placed at these points, the displacements along the orthogonal direction would cause the fiber to be placed in tension, and thereby distributing the stresses from struts. Displacements are nearly zero at the2048point lying in center since they lie on the neutral axis if this truss arrangement is considered as beam.2D-analysis of the same truss-arrangement with fiber inclusions was conducted to test the effect of fiber. One basic assumption is that the bonding between the truss and fiber is considered to be perfect. The material properties of the truss, meshing scheme, loading and boundary conditions all remained the same as the previous study.

The material properties of the truss, meshing scheme, loading and boundary conditions all remained the same as the previous study. For the fiber material stainless steel wire type 302 (elastic modulus = 193 GPa [40]) was chosen with a standard 1/8 inch diameter. The fiber was modeled as a T2D2 (no bending) truss element with no-compression to deactivate, as fibers do not provide stiffness against compression loading. Figure 6 (a) shows the placement of fibers.The fibers were placed on the top layer to maintain the symmetry of structure for loading. The truss arrangement is considered to be a section of a larger structure and hence the fibers were located in the center as well to show that they would engage otherwise. It is important to note that the fibers connect the nodes of the cellular structure to fulfill the purpose of incorporating tensegrity behavior. Also, due to the pre-stressing, the fibers in the top layer would engage even if they are under compression.

Also, due to the pre-stressing, the fibers in the top layer would engage even if they are under compression. However, this paper does not concentrate on the effects of pre-stress but rather lays out the concept of fiber-arrangement such that tensegrity behavior is seen.(a)RollerPinnedFibersCenter(b)Figure 6. a) Fiber - included in truss, b) Deformed truss with fibersFigure 6 (b) represents the deformation of structure with fibers. The performance was measured in terms of displacement in direction of loading at the center of the structure and the stresses associated with it. Figures 7 (a) and (b) depict the displacement (Y-direction) and stress of the center and for parametric loading on the structures respectively. The slope of curve for truss with fiber is greater, implying the structure has become stiffer.2049(a)(b)Figure 7.

The slope of curve for truss with fiber is greater, implying the structure has become stiffer.2049(a)(b)Figure 7. (a) Load vs displacement at center, (b) Load vs stress at centerIt can be noted that the slope of the curve with fiber is much steeper. Stresses in truss with fibers decreased by 79 MPa. This improvement in strength can be attributed to the appropriate location of fibers, which have been engaged in tension to reduce stress. The displacement was decreased by 0.43 mm by fibers making the structures stiffer. The effectiveness of fibers included in structures was convincing to place them in octet cellular structures using the same methodology.3.3 Octet- Cellular Structure ModelingAs proved in 2D truss-analysis the fiber inclusions do make the structures stronger and stiffer.

3D Beam analysis was conducted on octet structures to understand the physics of structures and gain insight into the placement of fibers.A 3D Beam structure was built to represent octet geometry as shown in Fig. 8 (a). The strut dimensions were based on the same slenderness ratio as selected in Section 3.1 Nine octet cells were repeated in length and two in width to form a structure suitable for 4-point bending test. The overall dimensions were 318.20 mm x 70.71 mm x 35.35 mm. The structure was assigned Al-6061 T6 material with B-31 mesh elements. This was specifically done to confirm the magnitude of bending stress present in the structure. The loading and boundary conditions for 4-point bending test were assigned as shown in Fig. 11 (a).Loading pointsRoller(a)(b)Pinned2050Figure 8. (a) Octet cellular with loading and boundary conditions, (b) Deformed octet structureThe deformation of the structure was studied as shown in Fig. 15 (c).

(a) Octet cellular with loading and boundary conditions, (b) Deformed octet structureThe deformation of the structure was studied as shown in Fig. 15 (c). Fibers were included at points as chosen in Section 3.2. The fiber material remained the same as mentioned in Section 3.2.1. As depicted in Fig. 9 (a), 3D-beams octet structure was modeled with fibers. As in the 2D model (Section 3.2), the fibers were represented as T3D2 no-compression truss elements since they have no stiffness against compression loads and cannot experience bending loads. Similar loading and boundary conditions were provided as that shown in Fig. 9 (b).Loading pointsRoller(a)(b)PinnedFigure 9. (a) Octet structure with fibers-loading and boundary conditions, (b) Deformed structure3.3.1 ResultsFigure 9 (b) shows the deformations occurring in octet structure with steel wires. The deformations in both the octet structures were studied in terms of stresses and displacements.

The deformations in both the octet structures were studied in terms of stresses and displacements. Figure 10 (a) and (b) represent the displacement at the loading points and maximum stress observed in structures for parametric loading respectively.(a)(b)Figure 10. (a) Load vs displacement at center, (b) Load vs stress at centerTable 1 provides the details of effectiveness of steel wires through displacements and normal stresses due to the presence of steel wires. The stiffness of the structure with steel wire2051increased through reduction in vertical displacement by 2 mm. Increase in strength due steel wires was about 178.3 MPa, which underlines the effect of tensegrity behavior.Table 1. Effectiveness of fibersOctet Structure Octet Structure with steel wiresMaximum Displacement (mm) 4.4 2.4Maximum Stress (MPa) 253.4 75.13.4 Tensegrity ObservationThe bending stress in both the structures was studied as shown in Fig. 11 (a).

11 (a). The tensionexperienced in all bottom wires were studied, as shown in Fig. 11 (b).(a)(b)Figure 11. (a) Bending stress in octet structure with wires (b) Tension in steel wiresThere is reduction of these stresses in octet structure with wires. This is attributed to the interplay between octet cellular shape and the tension in steel wires. This interplay is a consequence of tensegrity behavior where the tension in wires can effectively engage to reduce the stress due to the shape of the cell and wire arrangement.As observed in there is presence of bending stress in octet structure with wires, but the value is very low. The structure does not experience much of bending stress due to the octet arrangement. The tension in steel is seen to be increasing with the load. The structure with wires is shown to be stiffer and stronger through results in Table 1. This behavior is typical of tensegrity structures.

The structure with wires is shown to be stiffer and stronger through results in Table 1. This behavior is typical of tensegrity structures. Due to the presence of some small bending stress it is stated to be nearly tensegrity behavior.4 Experimental Test and Validation4.1 FabricationA manufacturing process to produce complex cellular metal composites using Binder Jetting AM process was established [41]. The process makes use of sand molds printed through Binder Jetting technology and metal is directly cast into them to produce the final part. In this paper the aforementioned process was adapted to produce cellular structures with fiber- reinforcements as shown in Fig. 12.2052The mold arrangement was split such that fiber embedding could be achieved. The mold- designs were created through a Boolean subtraction of Octet design from a block of exact overall dimensions. Molds were printed using the CAD-designs using Binder Jetting AM technology.

The mold- designs were created through a Boolean subtraction of Octet design from a block of exact overall dimensions. Molds were printed using the CAD-designs using Binder Jetting AM technology. The wires were embedded in the molds and the mold-setup was complete with additional packing of sand around it to secure the positions of wires. For this mold, A356 alloy was poured, which is a common aluminum alloy (Al-7wt%Si-0.3wt%Mg) [42] with a good combination of mechanical properties and castability. The temperature of metal in furnace was limited to 754.44 °C (1390 °F). The temperature while pouring, although not measured, is expected to be ~10-20 °C lower than the furnace. As shown in Fig. 12, the structure was successfully embedded with steel wires to attain a final part resembling the structure in Fig. 9.

As shown in Fig. 12, the structure was successfully embedded with steel wires to attain a final part resembling the structure in Fig. 9. Another structure without wires was fabricated alongside to complete the comparison study while testing.Mold Generation3D CADMold GenerationSplit MoldsMold SetupPrinted MoldsEmbed wireMold-SetupCastingCast PartMachined structureFigure 12. Process to create cellular structures with tensegrity behaviorThe weights of both structures were measured after machining with fiber-reinforced octet structure at 1.44 kg while the one without wires was 1.16 kg.4.2 TestingFour-point bending test was carried out on both the fabricated structures using a testing fixture built for this. The test was conducted on MTS Insight using the control module provided2053by MTS. Figure 13 (a) shows the structures set on the fixtures for testing. The region between loading points were subjected to constant bending moment and zero shear force.

Figure 13 (a) shows the structures set on the fixtures for testing. The region between loading points were subjected to constant bending moment and zero shear force. The maximum value of bending moment is to be found in this region. Tests were conducted at constant displacement rate of 2 mm/min to achieve quasi-static loading condition according to the military standards: MIL-STD-401 DIN 53291.The displacement of loading points and the loading data were acquired during the tests. Figure 13 (b) indicates the maximum displacements caused in both the structures during the test for the acquired loads.(a)(c)(b)Figure 13. (a) Four point bending test setup (b) Comparison of load and deflection in octet structures with and without steel wiresThe load needed to cause failure in octet structure with steel wires is observed to be 36.59 kN, which is 8.348 kN more than the failure load for octet structure without wires.

The slope of2054the curves (difference in slope - 1100) in elastic region as shown in Fig. 13 (b) suggest that the bending stiffness of octet structure with steel wires is larger, proving the structure for better performance.As shown in Fig. 13 (c) the failure first occurs in aluminum strut and the steel wires are still intact. The steel wires are still engaged and this increases the ultimate strength of the structure. The wires are seen to perform more efficiently in the plastic region since they were not pre-stressed initially but only held in enough tension to engage while testing. In the plastic region the wires are at their maximum potential and hence the ultimate strength is observed to have increased more than the stiffness.Figure 14. Comparison – Simulation and ExperimentFigure 14 depicts the differences between the simulation and experimental testing in the elastic region. For octet structure with wires the discrepancy is less than 3 %.

For octet structure with wires the discrepancy is less than 3 %. This error can be due to the non-uniformity of struts in the fabricated parts. While the struts were designed to be of diameter 7 mm, this wasn’t achieved. The average strut diameter was found to be around 7.3 mm, which makes the structure stiffer than the simulation. At lower strains, for the structure with wires the error is around 6 % while at higher strains it is around 15 %. This discrepancy can be attributed to the bonding issues between the wires and aluminum. The simulation perfect bonding while in reality this may not be achieved.5 Closure and Future Work2-D finite element analysis of 4-point bending test was conducted on a truss-arrangement to understand the mechanics and establish a method for the placement of fibers. The bending stresses experienced by octet structures, while wires engaged in tension to reduce stresses2055indicating nearly tensegrity behavior was achieved.

The bending stresses experienced by octet structures, while wires engaged in tension to reduce stresses2055indicating nearly tensegrity behavior was achieved. 3-D beam analysis of octet structures with fibers showed improvements in strength and stiffness due to incorporation of tensegrity behavior.The fibers effectively reduced the stress from adjacent sections by engaging in tension. The structures without wires experienced normal stress of about 253.4 MPa while the structures with wires reported stress of about 75.1 MPa. In the analysis it was seen that octet structures with fibers experienced very little bending stress (about 0.25 MPa). Nearly tensegrity behavior was achieved to improve strength and stiffness during 4-point bending test as discussed in Section 3.4. The octet structures were successfully fabricated through process chain as indicated in Fig. 12.

The octet structures were successfully fabricated through process chain as indicated in Fig. 12. Fiber- embedding was successfully accomplished in fabrication through molds created by using Binder Jetting AM technology.For an increase in weight of 0.28 kg, the octet structures with steel wires demonstrated higher stiffness and strength by failing at 36.59 kN. The same increase in strength can be achieved with even less weight-increase if fibers such as carbon-fiber, silicon-carbide and alumina are used since their stiffness to mass ratio is considerably higher than steel wire.In the future, a 3D solid model will be constructed for finite element analysis and validate experiments by simulating the 4 point bending test. The effects of fiber would be studied thoroughly to establish a criteria of design. A study on pre-stressing would help in increasing the overall stiffness of the structure and ultimate strength to a greater extent than what is observed in this study.

A study on pre-stressing would help in increasing the overall stiffness of the structure and ultimate strength to a greater extent than what is observed in this study. Material characterization will be carried out to understand the bonding between steel wire and aluminum.6 AcknowledgementsThis material is based upon work supported by the National Science Foundation under Grant No. 1462089. Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the views of the National Science Foundation. The authors wish to thank Ex-One for 3D printing molds. Also, the authors thank Erin Connelly and Patrick O’Brien for their help during fabrication and testing.References[1] V. S. Deshpande, N. A. Fleck, and M. F. Ashby, “Effective properties of the octet-truss lattice material,” Journal of the Mechanics and Physics of Solids, vol. 49, pp. 1747–1769, 2001.[2] M. F. Gibson, J.

SIGN OF PASSIVE DYNAMIC WALKING ROBOTS FOR ADDITIVE MANUFACTUREF. Modica*, F. R. Stöckli*, K. Shea*Engineering Design and Computing Laboratory, Department of Mechanical and Process Engineering, ETH Zürich, SwitzerlandAbstractOngoing research in the direction of printable, non-assembly mechatronic systems give rise to the need for multi-material printing, including electronics. However, there are robotic systems that do not use electronic components and still exhibit complex dynamic behavior. Such passive dynamic systems have the potential to save energy and component cost in the field of robotics compared to actuated systems. Ongoing research in computational design synthesis of passive dynamic systems aims at automatically generating robotic configurations based on a given task. However, an automated design-to-fabrication process also requires a flexible fabrication method.

However, an automated design-to-fabrication process also requires a flexible fabrication method. Towards the goal of printing functional, non-assembly passive dynamic robots using Fused Deposition Modeling (FDM), this paper explores designing and fabricating passive walking robots and all necessary components using single material FDM. Two configurations of passive dynamic walkers are re-designed and fabricated in this paper. For one of them all components are printed in one job and only little assembly after printing is needed. However, the gait cycle of the second configuration is much more sensitive to small parametric changes and therefore more flexible prototyping is needed in order to allow adjusting of the robot after printing. Moreover, FDM printed robotic joints with sufficient smoothness and axial stiffness are required and a variety of different joint assemblies are designed and tested for the robot prototypes.

Moreover, FDM printed robotic joints with sufficient smoothness and axial stiffness are required and a variety of different joint assemblies are designed and tested for the robot prototypes. Even though the most stable gait for the second robot is achieved using a metal bearing instead of the FDM printed ones, this is not necessary for the first robot example. The approach to prototyping with FDM presented in this paper allows achieving functionality through design iteration without incurring significant cost. To arrive at feasible solutions, a modular design approach allows to combine different joints, legs, feet and balancing weights and the connection points of the different elements are adjustable after printing, which makes it possible to shift the center of gravity and other variables of the robot.IntroductionAdditive Manufacturing (AM) can reduce post-processing cost by fabricating assemblies as monolithic parts in single processes.

However, the technology to fabricate functional components, such as an actuated robotic system with integrated electronics, is still in its infancy. In contrast, passive dynamic systems do not require such components and are therefore suited for single-material AM processes. A passive walking robot, for instance, walks down a slope with gravitational force alone and without actuators. The design of a passive walking robot helps to understand the mechanics of human walking and it allows actuated, walking robots to be built with greater energy efficiency. First attempts to build passive dynamic walking robots started in 1990 [1]. Passive dynamic systems have, in general, the potential to save energy and component cost in the field of robotics compared to actuated systems. Ongoing research in computationaldesign synthesis of passive dynamic systems aims at automatically generating robotic configurations based on a given task [2].

Ongoing research in computationaldesign synthesis of passive dynamic systems aims at automatically generating robotic configurations based on a given task [2]. To extend this approach to an automated design-to- fabrication process choosing AM as fabrication method can offer a lot of flexibility.The goal of this paper is to explore the potential of low cost FDM fabrication of passive dynamic walking robots. Based on existing passive walking robots, two suited configurations are chosen and redesigned to be fabricated using AM. Joints are the most complex parts and different joint designs are tested and compared. The passive walkers require joints with low friction as well as high axial stiffness. The more advanced of the two robots is built in a modular way, such that the single components can be replaced with different versions.

The more advanced of the two robots is built in a modular way, such that the single components can be replaced with different versions. In order to test the suitability of the parts, a variety of prototypes covering all components are built and the influence of the components on the gait of the passive walker are examined with the help of experiments. To achieve a stable gait, significant fine tuning is necessary, since this passive walker easily looses balance. A second, more robust configuration of a passive walker is built, which has a stable gait without requiring much balancing. The imprecision of 3D-printed components are less problematic for this configuration, due to less sensitive dynamics.BackgroundThe theory of passive walking has been studied extensively in literature. A mechanical model for passive dynamic walking can be described with the rimless wheel model [3] [4] and the more complex compass gait model [5] [6].

A mechanical model for passive dynamic walking can be described with the rimless wheel model [3] [4] and the more complex compass gait model [5] [6]. Analyses showing which parameters influence the gait of the passive walker and how they can affect the dynamics of the gait can be found in [7].The most intuitive and straightforward passive walker is a configuration with two legs and no knees as shown in Figure 1a. This robot constructed by Tedrake is made from wooden and metal parts with a pin joint as the hip [6].abcFigure 1: Passive Walkers. a: Tedrake [6]. b: Colman and Ruina [8]. c: McGeer [1].Pushing the passive walker sideways lifts the opposite foot off the ground and thereby solves the problem of foot clearance. When placed on a slope, this opposite foot will swing forward due to gravity until it touches the ground again. The energy dissipation through friction and collision are overcome by gravitational forces.

The energy dissipation through friction and collision are overcome by gravitational forces. With experiments, Tedrake determined which curvature of the feet provides a stable gait on a small decline.Colman and Ruina [8] built a passive walking robot with components from the Tinkertoy construction set, see Figure 1b. Contrary to Tedrake's construction, the Tinkertoy passive walker cannot stand. The Tinkertoy prototype exhibited a stable gait, although this is not achieved in simulations in [8] . Coleman and Ruina achieved a stable gait by changing the centers of mass of the legs.The third passive walker shown in Figure 1c was developed by McGeer [1]. A knee joint is included in this prototype to solve the problem of foot clearance. The robots presented above solve this problem by tilting sideways. McGeer’s robot demands higher quality of the components due to its greater sensitivity and complexity.A fourth configuration of a passive locomotion robot is shown in Figure 2.

McGeer’s robot demands higher quality of the components due to its greater sensitivity and complexity.A fourth configuration of a passive locomotion robot is shown in Figure 2. The kangaroo is able to “hop” down a slope without actuators. Similarly to McGeer's passive walker, the kangaroo has two permanently fixed outer legs. In between these legs, the kangaroo touches the ground with its upper limbs and tail. This causes a forward / backward motion of the body. While balanced on the upper limbs, the legs move forward and the cycle continues. The kangaroo has four contact points with the ground and is statically stable.Figure 2: Hopping Walker has a robust gait [9].Single-material AM has been used to fabricate non-assembly mechanical joints [10] as well as complete articulated toy characters [11]. Actuated deformable characters have been built using multi-material AM [12]. In [13] robots are automatically designed and the robot skeleton is fabricated using AM.

Actuated deformable characters have been built using multi-material AM [12]. In [13] robots are automatically designed and the robot skeleton is fabricated using AM. With these approaches, however, additional external actuators are required. To circumvent the need for the addition of electric parts, this paper demonstrates how low cost single-material FDM can be used to fabricate functional passive dynamic walking robots that do not require actuators. It investigates the possibility to use FDM fabricated parts only. In order to do so, extensive experimental testing and fine tuning of the printed walking robots is required. A variety of joints that can be printed with FDM are designed and tested, as well as compared with metal ball bearings.Concept DesignFrom the existing passive walking robots presented, two suitable configurations are chosen for fabrication.

The kangaroo is chosen because of its robust gate and the design from Colman and Ruina is chosen because of its low complexity and the configuration can be fine- tuned after printing by shifting the center of gravity to achieve a stable gait. While different feet curvatures, as used by Tedrake, could be easily fabricated using AM, they are not adjustable once printed and therefore Tedrakes configuration is not chosen.The design presented in this paper is shown in Figure 3a. It consists of the feet, the legs, a hip joint, an extension element that allows shifting the center of mass in all three dimensions and balancing weights with different masses. Connection elements allow to fix and unfix the components, as well as shifting the connection points.Figure 3b shows the geometric parameters that can be varied. 𝑏1, 𝑏2 and ℎ2 are the positions of the balance weights, ℎ1is the height of the hip joint and 𝑤 is the distance between the legs.

𝑏1, 𝑏2 and ℎ2 are the positions of the balance weights, ℎ1is the height of the hip joint and 𝑤 is the distance between the legs. Other important parameters are the curvature of the feet and the surface adhesion, as well as all the masses and moments of inertia of the feet, legs, connection elements and balance weights.abFigure 3: General ConfigurationIn addition to the passive walker presented in this section, which has only two contact points to the ground, a version of the kangaroo, shown in Figure 2, is redesigned for additive manufacturing. This configuration has four points that can be in contact with the ground and has a much more robust gate.Component DesignIn this section the components of the passive walker and their advantages and disadvantages in regard to the dynamics are presented. The components are designed and modelled using NX CAD software [14] and fabricated using a StrataSYS uPrint SE Plus printer.

The components are designed and modelled using NX CAD software [14] and fabricated using a StrataSYS uPrint SE Plus printer. This FDM machine uses an ABSplus thermoplastic material and prints with a layer thickness of 0.254 mm [15].Hip JointFor the hip joint, different designs are tested. The main challenge for the 3D printed components is to achieve a constant low friction during rotation, which is obtained by increasing the clearance between the moving parts of the joint. On the other hand, a smaller clearance is desired for more axial rigidity.a Figure 4: Bearing III. a: CAD model front- and back-view. b: Printed partbThe first design is shown in Figure 4. This bearing exhibits a smooth spinning behavior with constant friction, but significant play on the center axis of the hip joint. Secondly, a planetary gear is considered (see Figure 5).

This bearing exhibits a smooth spinning behavior with constant friction, but significant play on the center axis of the hip joint. Secondly, a planetary gear is considered (see Figure 5). Connecting one leg to the outer gear and the other leg to the inner gear results in a slight asymmetry in the moment of inertia of each leg. Moreover, the rigidity with respect to the center axis did not improve compared to the first bearing. Both bearings are printed with a range of different clearances between ball and casing and betweengears of the planetary bearing, in order to find a balance between rotational smoothness and axial stiffness.a Figure 5: Planetary Gear I. a: CAD model front- and back-view. b: Printed partbFigure 6 shows an approach to reduce the axial play in the joint. A pin connects both inner gears to one leg, whereas the other leg is connected to the outer gears (see Figure 6b).

A pin connects both inner gears to one leg, whereas the other leg is connected to the outer gears (see Figure 6b). While the rigidity did improve significantly, the asymmetry is more distinctive and the whole bearing is much heavier.abcFigure 6: Planetary Gear III. a: CAD model. b: Schematic diagram. c: Printed partIn order to create a stiff and yet smoothly running hip joint, a casing for a metal ball bearing is fabricated (see Figure 7). The casing allows a firm yet removable connection to the metal bearing. This configuration results in a sufficiently rigid joint.Table 1 summarizes the relevant properties of the different bearing configurations. The classification in Table 1 is made by a direct comparison of the different version of the hip joint regarding the labeled properties. Using the joints from Figure 4 and 5, one leg is connected to the outer gear and the other leg is connected to the inner gear.

Using the joints from Figure 4 and 5, one leg is connected to the outer gear and the other leg is connected to the inner gear. This causes a certain asymmetry, because the inertia of the inner and outer gear are different, which results in different dynamics for each leg. A symmetric behavior is beneficial, because it leads to a straight gait.abFigure 7: Casing for Metal Bearing. a: Setup. b: 608zz Ball BearingHip JointSmooth & constant frictionRigidity of center axisSymmetry WeightAM onlyFeasible for AMBearing I Bearing II Bearing III Planetary Gear I Planetary Gear II Planetary Gear III Insert Element0 0 0 - - 0 +0 0 0 0 - + +0 0 0 + - +0 + + 0 - 0+ + + + + + --* + + -* -* + +Table 1: Overview Hip Joints, Legend: + Good performance, 0 Average performance, - Poor performance *could not dissolve support materialLegs and FeetTo connect the hip joint to the legs several connection methods are considered.

Figure 8a shows a pin connection, where a holding device connected to the hip joint can be pinned to the leg. The connection is robust and allows for straight forward fixing and unfixing.abcdFigure 8: Legs, Feet and Connection. a: Pin Connection. b: Screw Connection, c: Leg, d: Connection Feet – LegTo obtain a pin connection where the pin can be inserted in the hole with the right amount of friction, the orientation of the component in the printer plays an important role. If the orientation of the hole in the printer is as in Figure 9 on the left, the diameter of the holes has to be increased by 0.05 mm. Placing the hole as in Figure 9 on the right fabricates a smoother edge of the hole.Figure 9: Orientation of Components in the PrinterTo build a lighter connection, a screw connection is designed (see Figure 8b). Figure 8a and 8b show the two leg designs with their different connection types.

Figure 8a and 8b show the two leg designs with their different connection types. The pin connection requires holes perpendicular to the center axis of the hip joint, whereas parallel holes are required for the screw connection. The leg in Figure 8c shows a configuration that allows use of both connection types. The pin connection can also be used to connect the feet to the legs (see Figure 8d).Extension Elements and WeightsAs described earlier, Colman and Ruina’s Walker (Figure 1) requires an element allowing to shift the center of mass and influence the moment of inertia. Figure 10 shows a design for the extension element on which balancing weights can be attached.Figure 10: Extension ElementThe weights and extension elements are connected by two pins that can be plugged into their counterpart. Different approaches have been tried varying the size of the holes and pins. Bigger holes simplify the fixing and unfixing of the weights and are less prone to breaking.

Different approaches have been tried varying the size of the holes and pins. Bigger holes simplify the fixing and unfixing of the weights and are less prone to breaking. On the other hand the shifting increment rises with bigger holes and therefore reduces the accuracy in the shift of the center of mass. The extension element can be fixed on the leg using the pin or screw connection presented in the previous section. Moreover, those connections allow shifting the center of mass along the z-axis (see Figure 3 for coordinate system).Experiments and ResultsThe experiments are conducted on a ramp of varying slope shown in Figure 11. Two rubber stickers with high adherence to prevent slipping are attached on the top of the ramp. A stable gait for Ruina’s passive walker was not found in simulation, but only found by experimenting with a physical model in [8]. Therefore it is crucial to be able to shift the balancing weights after printing to experimentally find a stable gait.

Therefore it is crucial to be able to shift the balancing weights after printing to experimentally find a stable gait. Another reason for the need for adjustability is the inaccuracies introduced by the FDM process. The initial state of the passive walker is induced by hand. Several iterations show that it is beneficial to tilt the walker towards the stance leg, which is in contact with the ground, and to rotate the swing leg backward. This proves to be an initial state, after which several stable steps are likely to follow.Figure 11: Experimental SetupThe first prototype shown in Figure 12a consists of a 3D printed bearing as shown in Figure 4 as a hip joint. Both the hip joint and the extension element are fixed on the leg with pin connections and are therefore adjustable.ab Figure 12 a: Prototype I, b: Prototype IIWhile conducting experiments with Prototype I, the following complications are observed. It tends to fall forward or backward, due to the high center of mass.

It tends to fall forward or backward, due to the high center of mass. The shifting increments of the balance weights are too large. The bearing is not stiff enough, meaning the center axis tilts easily. Using the planetary gear as a hip joint did not show a considerable improvement in the gait.To lower the center of mass for Prototype II (see Figure 12b) heavier feet are used and the hip joint is replaced by a metal bearing. Using a metal bearing also reduces the play in the center axis. Designing an extension element with smaller holes reduces the shifting increments of the balancing weights.Conducting experiments with Prototype II shows a stable gait for about nine steps. This means that the axial stiffness of the metal bearing as a hip joint is essential for a stable gait, which could not be achieved with the FDM produced bearings. In Figure 13, the mechanical properties of the left half of Prototype II are given.

In Figure 13, the mechanical properties of the left half of Prototype II are given. The walking direction is in positive x- direction.𝐶𝑀𝑥, 𝐶𝑀𝑦, 𝐶𝑀𝑧 [𝑚𝑚] 𝐼𝑥, 𝐼𝑦, 𝐼𝑧 [𝑘𝑔 𝑚𝑚2] 𝐼𝑦𝑧, 𝐼𝑥𝑧, 𝐼𝑥𝑦 [𝑘𝑔 𝑚𝑚2] 𝑅𝐹𝑜𝑜𝑡[𝑚𝑚] ℎ1 [𝑚𝑚] 𝑤 [𝑚𝑚] 𝑚 [𝑘𝑔]3.4, 88.1, 48.8 1627, 437, 1260 401, -17, -66 27 126 40.5 0.111Figure 13: Parameters of Prototype II; All values for the left half of prototypeabFigure 14: Kangaroo, a: Picture, b:SetupThe kangaroo, as shown in Figure 14, has four points that can be in contact with the ground, which makes it easier to achieve a stable gait as little balancing is needed. From a template [9] the form is adopted and the Planetary Gear I is used as the hip joint. Two pins connect both legs to the body of the kangaroo.

From a template [9] the form is adopted and the Planetary Gear I is used as the hip joint. Two pins connect both legs to the body of the kangaroo. The experiments using this passive walker shows that a stable gait is possible with this configuration, which is completely printed using FDM.DiscussionObtaining a stable gait with a FDM printed passive walker with a similar configuration to the Tinkertoy design is hard to achieve. The stable gait depends on many factors. The hip joint has to exhibit a certain rotational smoothness that has to remain constant over one step. With 3D- printed hip joints, this property is hard to obtain while keeping the axis rigid. A more rigid hip joint reduces the smoothness and vice versa. A suitable compromise could not be found in this project. In addition to the smoothness and rigidity, the hip joint has to exhibit symmetric behavior resulting in identical dynamics for both legs.

A suitable compromise could not be found in this project. In addition to the smoothness and rigidity, the hip joint has to exhibit symmetric behavior resulting in identical dynamics for both legs. Using a metal bearing eliminated the problems and showed an improvement in the walking behavior.The kangaroo (see Figure 14) does not require much balancing in order to achieve a stable gait due to its configuration. Without this requirement, the FDM planetary gear allows a stable gait, even with the known imprecision. The kangaroo shows that stable passive dynamic locomotion is achievable using only FDM printed components.Both configurations can be printed in one job. However, they are printed in parts and fully assembled after printing. When printed as an assembled prototype, there is support material added in the bearing. After removal of the support material the bearing surface is rough. The bearing of the kangaroo in Figure 14 is designed to be printed without support material.

After removal of the support material the bearing surface is rough. The bearing of the kangaroo in Figure 14 is designed to be printed without support material. However, if the whole kangaroo is printed in one job with the legs already attached, the bearing would be filled with support material and would not operate as smoothly after support material removal. However, it is expected that it will be possible in future work to print a fully assembled working model by adjusting the layout of the assembly.ConclusionTwo different configurations of existing passive dynamic locomotion robots, a Kangaroo and a Tinkertoy design, are redesigned to be fabricated using low cost FDM. A fully working prototype of the Kangaroo configuration with a more robust gait is printed in one job and tested successfully thus demonstrating a passive dynamic locomotion robot fabricated only out of FDM parts.

For the second Tinkertoy configuration, which has a more sensitive gait, different types of FDM printed joints are designed and tested. However, a stable walk is obtained only by adding a metal bearing to the FDM parts. Moreover, a modular structure of the robot is used, allowing the exchange of parts as well as changing system parameters after printing, for example the center of gravity, which is crucial in the experimental testing and fine tuning of this type of passive walker. Different connection elements and feet are designed and tested. Through the success of the Kangaroo with an FDM printed planetary gear, a new approach to 3D printing of passive dynamic robotic systems is shown.References[1] T. McGeer, "Passive Dynamic Walking," the international journal of robotics research, vol.9.2, pp. 62-82, 1990.[2] F. R. Stöckli and K.

McGeer, "Passive Dynamic Walking," the international journal of robotics research, vol.9.2, pp. 62-82, 1990.[2] F. R. Stöckli and K. Shea, "A Simulation-Driven Graph Grammar Method for theAutomated Synthesis of Passive Dynamic Brachiating Robots," in ASME International Design Engineering Technical Conferences, Boston, 2015.[3] M. J. Coleman, A. Chatterjee and A. Ruina, "Motions of a rimless spoked wheel: a simple three-dimensional system with impacts," Dynamics and Stability of Systems, vol. 12.3, pp. 139-159, 1997.[4] H. Chen, "Passive dynamic walking with knees: A point foot model," Diss. MassachusettsInstitute of Technology, pp. 15-20, 2007.[5] A. Goswami, B. Espiau and A. Keramane, "Limit cycles and their stability in a passivebipedal gait," in Robotics and Automation, IEEE International Conference on. Vol. 1. , 1996.[6] R. L. Tedrake, " Applied Optimal Control for Dynamically Stable Legged Locomotion,"Diss. Massachusetts Institute of Technology, pp.

SIGN OF PERIODIC CELLULAR STRUCTURES FOR HEAT EXCHANGER APPLICATIONSVikas Kumar1, Guhaprasanna Manogharan1, Denis R. Cormier21 Department of Industrial & Systems Engineering North Carolina State University, Raleigh, NC 276952 Department of Industrial & Systems Engineering Rochester Institute of Technology, Rochester, NY 14653AbstractPowder based metal additive manufacturing processes generally produce parts with a textured surface. Although surface roughness is undesirable in most cases, it can be advantageous for applications such as heat exchangers or catalyst support structures. While stochastic metal foams have been used for these applications, it is not possible to tailor the foam geometry to the application. SFF techniques permit designers to use different cell geometries and orientations to achieve specific performance objectives.

SFF techniques permit designers to use different cell geometries and orientations to achieve specific performance objectives. More specifically, the unit cell type and size, ligament size, and cell orientation have a considerable influence on surface area, volume, density and fluid flow behavior of the lattice structure. This paper illustrates the effect of the orientation of hexagonal periodic cellular structures on heat transfer and pressure drop. Finite element analysis of air flowing through the hot hexagonal periodic cellular structures shows the effect of orientation on heat transfer and pressure drop. The analysis indicates that the ideal orientation is dependent on the specific application.IntroductionPeriodic cellular materials are a class of materials whose properties can be tailored to enhance thermal, electrical and/or mechanical properties.

Open celled materials with high porosity, large surface area and light weight are very well suited for heat exchange applications. Cellular materials have historically been classified as either stochastic (e.g. random cell geometry) or periodic (e.g. engineered repeating cell geometry). There are several methods of fabricating periodic cellular structures described in the literature, but Solid Freeform Fabrication (SFF) techniques give maximum flexibility in terms of tailoring the shape and size of the lattice to the needs of the application. SFF techniques are capable of producing near net shaped parts with textured surfaces as shown in Figure 1. Although surface roughness is undesirable in many applications, it can be advantageous for applications such as heat exchangers or catalyst support structures. Surface roughness can affect both the heat transfer and pressure drop characteristics of the fluid flow through periodic cellular structures.

Surface roughness can affect both the heat transfer and pressure drop characteristics of the fluid flow through periodic cellular structures. Heat transfer and pressure drop is also affected by cell geometry, cell size, foam density and the cell orientation, where cell size and foam density are a function of ligament radius and ligament length for a given cell geometry in periodic cellular structures.A great deal of research has been done with regard to estimating and optimizing the foam density and cell geometry for heat transfer applications, however there has been relatively little work on optimization of the cellular structure’s orientation with respect to fluid or gas flow. With SFF techniques, virtually any cell geometry at any orientation and of any ligament radius can be fabricated.

With SFF techniques, virtually any cell geometry at any orientation and of any ligament radius can be fabricated. The Electron Beam Melting (EBM) process is one SFF technique which has the capability to produce fully dense metal parts in materials of interest for heat exchanger applications such as aluminum or copper (Figure 2).(a) (b)Figure 1 Micrographs of Ti-6Al-4V hexagonal cellular structure made by EBM process before (a) and after (b) sand blastingIn the present work, a hexagonal periodic cellular structure made via the EBM process is modeled, and the effect of orientation on heat transfer and pressure drop has been analyzed. A combination of five different lattice orientations (0, 30, 45, 60, and 90 degrees) in each of the X, Y, and Z axes has been considered, making a total 125 different lattice orientations.

A combination of five different lattice orientations (0, 30, 45, 60, and 90 degrees) in each of the X, Y, and Z axes has been considered, making a total 125 different lattice orientations. The hexagonal cellular structure (Figure 3a) has the same shape with respect to the X and Y axes, hence an orientation of 30 and 60 degrees in the X axis yields the same effective cell orientation with respect to gas flow direction (Figure 3b). When redundant cell orientations are excluded, the number of unique orientations is reduced to 15. Finite element analyses of those 15 orientations have been performed using COSMOS Floworks 2008 to estimate the heat transfer and pressure drop.

Finite element analyses of those 15 orientations have been performed using COSMOS Floworks 2008 to estimate the heat transfer and pressure drop. The attributes of the hexagonal cellular structure compared using three non-dimensional groups include (a) the Nusselt number (Nu) characterizing the heat transfer, (b) the friction factor (f) for the pressure drop and (c) the Reynolds number (Re) for the fluid flow rate.(a)(b)Figure2 (a) Ti-6All-4V and (b)) Al-6061 ceellular perioodic structuures made by EBM proa) (a(bb)Figurre 3 (a) Persspective vieww (b) First a periodic cangle orthog cellular strugraphic pro uctureojection of the hexagoncessalBackgroundTransport phenomenon in stochastic metal foams has received considerable attention during the last decade. Lee et al. [1] studied the application of metal foams as high performance air-cooled heat sinks in electronics packaging. Sathe et al.

Lee et al. [1] studied the application of metal foams as high performance air-cooled heat sinks in electronics packaging. Sathe et al. [2] studied combustion in metal foams for application in porous radiant burners. Hunt and Tien [3], Calmadi and Mahajan [4], Boomsma and Poulikakos [5], Lu et al. [6] and Fuller et al. [7] have addressed the thermal characteristics of metal foams. Wadley and Queheillalt [8] have compared the thermal properties of copper textile-based lattice structures with the stochastic open-cell metal foams.Tian et al. [9] studied the effects of topology upon fluid flow and heat transfer within cellular copper structures of different density. Experiments were performed with the prototype sandwiches of two different cell geometries and two orientations, 0 and 90 degrees with respect to the front plane. The friction factor was found to be independent of coolant velocity and dependent on open area ratio.

The friction factor was found to be independent of coolant velocity and dependent on open area ratio. For a particular porosity, the heat dissipation rate was found to be directly proportional to the surface area density. For a fixed surface area of pure copper, the optimal porosity was found to be 0.75. The overall heat transfer was primarily dependent upon porosity and surface area density. It was weakly dependent upon orientation. Lu et al. [10] reviewed the thermal characteristics of metallic sandwich structures by analytical modeling and numerical simulation. The Reynolds number, Nusselt number and friction factor of different types of sandwich structures were calculated and compared to conclude that topology preference is highly application dependent.Heat Transfer Model and CalculationA CAD model of a hexagonal periodic cellular structure (Figure 3a) of size 40mm x 40mm x 11.5mm was made on the SolidWorks 2008 CAD package.

The lattice was merged with a square tube such that it was possible to simulate gas entering the tube on one end (e.g. the inlet), flowing through the lattice, and then exiting via the outlet on the other end of the square tube. Although process parameter and post processing will significantly influence surface roughness, the cellular structure was assigned a surface roughness of 300 microns for purposes of this study. That figure is representative of a part fabricated with intentionally high surface roughness for purposes of heat exchange. The lattice heat exchanger was given a constant temperature of 120° C with an inlet gas temperature of 20° C flowing at a velocity of 9.55 m/s. Finite element analysis (FEA) of the heat transfer and pressure drop was conducted using COSMOS FloWorks 2008 for each lattice at orientation. The outputs of the FEA were average temperature and pressure of air at the outlet and the heat flux of the cellular structure perpendicular to the air flow.

The outputs of the FEA were average temperature and pressure of air at the outlet and the heat flux of the cellular structure perpendicular to the air flow. The surface area and the volume of material in the cellular structure were determined from the SolidWorks 2008 CAD model of the cellular structure.The temperature difference between the air at the inlet and outlet refers to the amount of heat transferred to the air by the cellular structure at a given orientation. Similarly, pressure drop reflects the amount of hindrance in the path of the fluid. Relative density of the lattice structure was obtained by calculating the ratio of the volume of the material in the structure and the totalrectangular bounding box volume occupied by it. To obtain the Nusselt number, Reynolds number and friction factor, the hydraulic diameter (Dh) must be calculated.

To obtain the Nusselt number, Reynolds number and friction factor, the hydraulic diameter (Dh) must be calculated. This is given by 1c l 4 cρ−D h=where cl is the cell edge length , ρ is relative density, and c is the cell geometry constant [11]. uN ) is defined as the ratio of convective to conductive heat transfer across The Nusselt number ( the boundary and is given byNwhere h is the local heat transfer coefficient andu=hD h k fk is the thermal conductivity of air. Reynold’sfnumber (eR ) defines the flow and is given byUDρ f h μ f where fμ is the viscosity of the fluid (air). The friction factor ( sf ) is a measure of the resistance provided to the flow by the cellular structure.

Orientation has been denoted in terms of the rotation angles and Z axes respectively. Figure 4 shows how the pressure drop is influenced by the lattice orientation with respect to the gas flow direction. For example, orientation #15 (45,30,45) produces the highest pressure drop while orientation #4 (0,45,0) produces the lowest pressure drop among the considered orientations. This is intuitively explained upon examination of Figures 5(a) and 5(b). Figures 5(a) and 5(b) show a view of the lattice from the inlet looking towards the outlet for orientation #4 and orientation #15 respectively. In orientation #4, the gas has a relatively unobstructed path from inlet to outlet, hence the pressure drop can reasonably be expected to be relatively low.

In orientation #4, the gas has a relatively unobstructed path from inlet to outlet, hence the pressure drop can reasonably be expected to be relatively low. In orientation #15, the gas has a tortuous path from inlet to outlet, hence the pressure drop can intuitively be expected to be much higher.,Figure 6 shows the predicted temperature drop from inlet to outlet as a function of lattice orientation. Orientation #4 (0,45,0) has the highest temperature difference while orientation #15 (30,0,0) has the lowest. The effect of orientation on heat transfer and pressure drop can further be explained in terms of the Nusselt number and friction factor. A Nusselt number of unity indicates sluggish flow in which heat transfer due to convection is equal to conduction. The Nusselt number seems to follow the same trend as the Reynold’s number for a given orientation. Higher value of friction factor shows greater hindrance to the path of the fluid.

Generally speaking, cellular materials the produce large pressure drops will require greater pumping power. the constraints are maximum allowable temperature in the metal and/or the fluid, or the pressure drop or pumping power. The thermal properties obtained in Table-1 can be used to design heat exchanger for required value heat transfer and pump capacity, implying that orientation preference depends on the application. Comparison of the performance of hexagonal periodic cellular structures with the thermal properties of stochastic and periodic structures [9,10] has been done in Figure 10. The hexagonal periodic cellular structure at the considered Reynolds number and density is close to the performance of the corrugated duct and louvered fins.

The hexagonal periodic cellular structure at the considered Reynolds number and density is close to the performance of the corrugated duct and louvered fins. The low friction factor implies lesser resistance to the fluid flow, and the lower nusselt number shows lower convection compared to copper textile mesh.Conclusions and Future WorkThis paper describes preliminary modeling effort targeted to the design of custom lattice structures that are tuned to the needs of a particular application. In this work, the influence of lattice orientation on flow and heat transfer properties has been modeled and analyzed. As expected, the orientation of the lattice with respect to gas flow direction does influence heat transfer efficiency and pressure drop.The next step is to model the influence of different types of unit cell geometries and different lattice densities.

One pressing issue for which no good solution has yet been devised is the issue of increased surface area in lattices fabricated via powder-based SFF processes. While FEA packages allow the user to specify surface roughness, the FEA does not actually use the surface roughness value to increase surface area. Rather, it is specified as a means of reflecting increased surface drag. However, surface area is known to have a very significant effect on heat transfer. It has been estimated that the surface area of an intentionally rough EBM latticestructure is approximately double the surface area computed by the CAD model with smooth struts.

It has been estimated that the surface area of an intentionally rough EBM latticestructure is approximately double the surface area computed by the CAD model with smooth struts. While the simulation here is useful for describing trends, better analytical methods are needed to more accurately reflect the increased surface area of the actual lattice structures.In order to determine just how significant the influence of surface area is on EBM- produced heat exchangers, efforts are currently underway to fabricate simple aluminum and copper lattice heat exchangers whose geometry matches those shown in Figure 3. Actual performance will be compared with predicted results, and those results will be used to successively refine and improve the analytical models presented in this paper as a starting point.References[1] Y. C. Lee, W. Zhang, H. Xie and R. L. Mahajan, 1993, ‘‘Cooling of a FCHIP Package With 100 W, 1 cm2 Chip,’’ Proceedings of the 1993 ASME Int. Elec. Packaging Conf., Vol.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesign of Spatially Varying Orientation Lattice Structures Using Triply Periodic Minimal SurfacesChongyi Wei, Douglas E Smith*Department of Mechanical Engineering, Baylor University, Waco, TX 76706AbstractInterest continues to grow for lattice structures produced by additive manufacturingmethods that are described by triply periodic minimal surface (TPMS). Tunable parameters that define the TPMS provide unique design flexibility where prior research has focused on designing hybrid or functionally graded TPMS structures. In this paper, a new strategy is proposed to include an orientation angle and volume fraction of each lattice cell simultaneously when defining structures derived from TPMS.

In this paper, a new strategy is proposed to include an orientation angle and volume fraction of each lattice cell simultaneously when defining structures derived from TPMS. The algorithm iteratively solves an underlying partial differential equation with the finite difference method to obtain a smooth, continuous lattice structure with a spatially varying orientation angle. The resulting lattice structure can be combined with other types of TPMS models using Gaussian radial basis and distance functions to achieve multi-TPMS lattice designs. The spatially varying lattice structures can also take the advantage of the directional effective modulus of TPMS to improve the strength the performance of lattice design.1.

The spatially varying lattice structures can also take the advantage of the directional effective modulus of TPMS to improve the strength the performance of lattice design.1. IntroductionThe mechanical and structural compatibility between orthopedic implant and human body is vital for successful implantation since a large mismatch in modulus can lead to stress shielding due to the uneven stress distribution at the bone-implant interface [1], [2]. Cellular lattice structures have been shown to be ideal for tailoring the properties of implants so as to avoid stress shielding. Within the wide range of lattice structures often considered, triply periodic minimal surface (TPMS) structures offer a unique combination of strength and biological compatibility [3]. The advancement in additive manufacturing has greatly facilitated the fabrication of lattices and thus allowed for the adaptation in the design process in biomaterials.

The advancement in additive manufacturing has greatly facilitated the fabrication of lattices and thus allowed for the adaptation in the design process in biomaterials. Therefore, lattice structures are increasingly employed in numerous applications, such as tissue engineering and other areas of mechanical engineering [4-6].Yoo [7] developed an efficient internal pore architecture design system based on a triplyperiodic minimal surface unit cell library and associated computational methods to assemble TPMS unit cells into an entire scaffold model. Due to the unique design flexibility of TPMS lattice structures, the lattice material properties can be easily tuned to produce a desired output property. When the unit cell porosity (defined as 1- ρ) and/or the unit cell size within a lattice is spatially varied to achieve a desired functionality, it is referred as functional grading [8].

When the unit cell porosity (defined as 1- ρ) and/or the unit cell size within a lattice is spatially varied to achieve a desired functionality, it is referred as functional grading [8]. In addition, a hybrid lattice structure can be obtained by combining two or more type of lattice structures together [9]. Further, the implicit surface mathematical function of TPMS lattice offers great advantage for modifying the cellular topology to generate anisotropic properties. However, there are limitations when tailoring the spatial structure and its properties when employing a1787single type of TPMS lattice. For example, as evident from the mathematic function, only one iso- surface parameter can be tuned to change the relative density.A variety of heterogeneous porous scaffold with gradients in porosity and pore size has been designed by using an efficient implicit interpolation method [10].

Numerical results show that the proposed scaffold design method has the potential benefits for accurately controlling the spatial porosity distribution within an arbitrarily shaped scaffold while maintaining the advantage of the TPMS-based unit cell. Afshar et al. [11] performed experimental and numerical investigations to determine the mechanical properties and deformation mechanisms of a linearly graded porosity scaffold. His work showed that the energy absorption for a uniform lattice is higher than those for graded structures and depends on loading direction. It is possible not only to change porosity through TPMS-based lattice structures, but pore architecture can be spatially varied to combine mechanical responses of different pore architecture.Cheikho et al. [12] presented an advanced and versatile method to design graded circularporous 2D structures based on the conformal mapping of unit cells.

[12] presented an advanced and versatile method to design graded circularporous 2D structures based on the conformal mapping of unit cells. His approach provided porous lattices obtained using a periodic transformation into a circular pattern by controlling the phase and amplitude of the periodic domain. The versatility allowed for tunable anisotropy by selecting the most appropriate structure for a given clinical application, depending on the desired porosity, pore size, and effective mechanical properties. Ma [13] proposed a design approach for heterogeneous porous scaffolds, which was obtained by discretizing the original model using the conformal refinement of the hexahedral mesh followed by a mapping of the TPMS unit onto existing mesh elements with shape functions.

The TPMS lattice generated on the uniform grid was mapped into the predefined mesh patterns such that the re-meshed scaffold could be designed to the appropriate elastic properties.To the best of our best knowledge, there is no studies that discuss TPMS lattice structures with both spatial varying infill volume fraction and orientation. Consequently, the first objective of the current study is to present an advanced framework for designing spatially varying TPMS lattices by changing the relative density and orientation for each unit lattice cell. Secondly, the design of heterogeneous lattice with spatially varying properties is demonstrated by two hybridization methods with irregular boundaries. Lastly, the versatility and limitations of the proposed method and its potential application to various tissue applications is discussed.2. Methodology2.1.

Lastly, the versatility and limitations of the proposed method and its potential application to various tissue applications is discussed.2. Methodology2.1. The design of spatially varying TPMS latticeTPMS lattice structures [14] are typically defined mathematically to have an approximately zero mean curvature at every point on the surface without self-intersection. A widely used TPMS form is the Schwarz’s Primitive-surface [15] as appearing in Fig. 1 (a), where the minimal surface is defined by the level set function(1) 𝐹(𝑥, 𝑦, 𝑧) = cos(𝑎𝑥) + cos(𝑎𝑦) + cos(𝑎𝑧) − 𝑡 > 0 where the parameter t controls the continuity of the cell surface which determines the volumetric solid bounded by the surface, and a defines the size of cubic unit cell. When 𝐹(𝑥, 𝑦, 𝑧) = 0, it1788defines a 3D boundary between positive F and negative F.

When 𝐹(𝑥, 𝑦, 𝑧) = 0, it1788defines a 3D boundary between positive F and negative F. Specifically, a candidate structure can be modeled using a 3D surface given by 𝐹(𝑥, 𝑦, 𝑧) − 𝑡 = 0 which defines a wide range of unit cell lattices with different volume fractions defined by t. Otherwise, the 𝐹(𝑥, 𝑦, 𝑧) can be modified to define other types of TPMS lattice structure, such as Gyroid and IWP shown as Fig. 1 (b) and (c) (not studied here).The level set function in Eq. (1) is limited to being able to the generation of uniformTPMS structures which only require the size of the unit cell size and the parameter t. The uniform TPMS has been extended to include a continuous sigmoid weight function to form hybrid types of TPMS structures with a functional dependance on the coordinate location [9]. These hybrid structure have a unit cell orientation that remains unchanged.

These hybrid structure have a unit cell orientation that remains unchanged. In addition, other research has shown that 3D periodic structures can be modeled by the interference to a four non- coplanar laser beams in the photonic crystals of physics field [16]. Here it is desirable to functionally grade and/or spatially vary the periodicity of the structure to form a metamaterial. The initial approach to develop a spatially varying lattice was adopted to guide the flow of the light in a preferred direction in prior work [17].Fig. 1 TPMS unit cell: (a) Primitive type unit cell; (b) Gyroid type unit cell; (c) IWP type unit cell. The shown TPMS structure with relative density 0.3, the t value is 0.7066, 0.6044 and 0.0049 respectively.To define a spatially orientation varying TPMS, the basic unit cell defined in Eq. (1)which describes a specific geometry is spatially varied to the form the final hybrid lattice.

(1)which describes a specific geometry is spatially varied to the form the final hybrid lattice. To this end, a periodic unit cell can be decomposed into a complex Fourier series where each term is spatially harmonic which can be interpreted as a one-dimensional sinusoidal grating that are combined to reconstruct the original unit cell. A finite set of spatial harmonics is truncated to construct the unit cell for practical implementation as(cid:3014) 𝜀(cid:3048)(cid:3030)(𝑠⃗) = (cid:3533) 𝛼(cid:3036)𝑒(cid:3037)(cid:3012)(cid:4652)(cid:4652)⃗(cid:3284)∙(cid:3046)⃗ (cid:3036)(cid:2880)(cid:2869)where 𝛼(cid:3036) represents the Fourier coefficient of the 𝑖(cid:3047)(cid:3035) spatial harmonic, obtained from the decomposition of the baseline unit cell. Otherwise, the specific expression can be reformulated as a generalized one as Eq. (2).

The single unit cell that describes the periodic structure that is to be spatially varied needs to be transformed into a set of planar gratings by Fourier transform, and the periodic array of elements are reduced into a Discrete Fourier Transform (DFT) which is then truncated to limit the number of terms. Therefore, the DFT is performed on the baseline unit cell to obtain each planar grating. Each of these resulting gratings are then spatially varied individually by including the orientation and accumulated to obtain the overall final lattice.1791(10)(11)(12)(13)(14)The spatially varying cell volume fraction over the lattice is incorporated following thesolution of the gradient equation governing the orientation as described above. There are two methods for mapping the relative volume fraction to the complete lattice structure.

There are two methods for mapping the relative volume fraction to the complete lattice structure. The first one is based on average density where the difference in relative density between adjacent cells is averaged has been shown lead to smooth geometric transition [22]. Secondly, the spatially variant fill fraction of the TPMS lattice structure can be implemented by a threshold technique to the preliminary binary distribution given as 𝜀(cid:2869), 𝜀(cid:2870),𝜀′(𝑠⃗) < 𝛾(𝑠⃗) 𝜀′(𝑠⃗) ≥ 𝛾(𝑠⃗)𝜀(𝑠⃗) = (cid:4682)where 𝛾(𝑠⃗) is the interpolated value from the previous calculation of the single unit cell relative volume fraction obtained by varying the iso-surface parameter t [17].

After the unit lattice cell are established from the threshold method, a linear interpolation is utilized to construct the relative density for the complete lattice structure.2.2 Designing heterogeneous TPMS lattice structureThe natural structure has specific porous structures in different regions according to theservice requirements, which leads to the heterogeneous structures [23].

Methods Implementation and Results3.1. Spatially varying TPMS lattice structureAs an example of constructing of spatially varying orientation TPMS lattice structure, theorientation distribution and relative density distribution field need to be defined firstly. The varying angle distribution can be generated by some mathematic function or the optimized results from topology optimization. After solving Eq. (7) numerically with Eq. (8), the grating phase 𝛷 can be incorporated into the exponential functions to further construct the 3D surface for TPMS lattice. The predefined orientation distribution and reconstructed spatially varying TPMS for gradually varying and self-defined varying field appear in Fig.2 and Fig.3 respectively.Fig. 2 The uniform gradual spatially varying Primitive TPMS lattice: (a) the varying lattice orientation and relative density distribution; (b) the reconstructed TPMS lattice.Fig.

2 The uniform gradual spatially varying Primitive TPMS lattice: (a) the varying lattice orientation and relative density distribution; (b) the reconstructed TPMS lattice.Fig. 3 The self-defined spatially varying Primitive TPMS lattice: (a) the varying lattice orientation and relative density distribution; (b) the reconstructed TPMS lattice.17943.2. Laminate or recursive TPMS lattice designYang [26] proposed a skewed TPMS lattice with a coordinate transformation to tune theanisotropy for a single TPMS cell rather than combining various TPMS types with weight coefficients.

Yang’s skew transformation can be viewed as a shift in the global coordinate of the cubic cell for TPMS lattice, the relationship is defined as𝐹(cid:2869) (cid:4672)𝑥′, 𝑦, 𝑧(cid:4673) > 0, where 𝑥′ = 𝑥 − (𝑧 + 1) tan 𝛼(cid:3051)(cid:3053)where 𝛼(cid:3051)(cid:3053) is the angle for the skew transformation, 𝐹(cid:2869) is skewed 3D iso-surface.In addition, a hybrid TPMS lattice with varying layer orientation can be generated bycombining a skewed TPMS lattice structure with a continuous sigmoid function as in Eq. (16), and the laminate (a concept borrowed from composite structure) lattice structure can be obtained by using the sigmoid function recursively shown as Fig. 4 (a). However, this method may lead to the structural discontinuity between two layers. The spatially varying lattice algorithm proposed in the current research can also be utilized to generate the laminate lattice with predefined orientation filed but better microstructure transition shown as Fig. 4 (b).Fig.

4 (b).Fig. 4 The laminate TPMS designed: (a) the laminate TPMS generated by hybrid with skew transformation; (b) the laminate TPMS generated by spatial varying transformation.3.3. Hybrid heterogeneous TPMS latticeA hybrid TPMS lattice can also be constructed using a weight function based on thesigmoid function in Eq. (16) where the hybrid transition area comes from the location function defined in Eq. (17).1795(26)Since GRBF is a real-valued function that depends on the distance between the input and some fixed points, the sum of radial function can be used to approximate the given boundary. An example of a computed heterogeneous TPMS with variant orientation for Primitive structures generated by GRBF appears in Fig. 5, where a uniform Gyroid and IWP lattice is also included that both have an irregular enclosed shape. Unfortunately, the lattice orientation leads to noncontinuous structure and defects on the transition boundary.

Unfortunately, the lattice orientation leads to noncontinuous structure and defects on the transition boundary. Therefore, the GRBF method needs to be refined to apply on spatially orientation varying TPMS lattice structures.Alternatively, complex line segments given by control points are represented by thedistance function in Eq. (25) are introduced. Here the distance function is incorporated into the sigmoid weight function to tune the transition area between different lattice types. The heterogeneous lattice with variant orientation angles for the Primitive type lattice is shown as Fig. 6 where the distance field function shows a good hybrid TPMS lattice structure compared to Gaussian Radian Basic Function representation.Fig. 5 Heterogeneous TPMS generated by GRBF: (a) the combination of P type and Gyroid type TPMS; (b) the combination of P type and IWP type TPMS.Fig.

5 Heterogeneous TPMS generated by GRBF: (a) the combination of P type and Gyroid type TPMS; (b) the combination of P type and IWP type TPMS.Fig. 6 Heterogeneous TPMS generated by distance field function: (a) the combination of P type and Gyroid type TPMS; (b) the combination of P type and IWP type TPMS.17964. Discussion and limitationThe spatial varying algorithm provides a significant advantage for accommodating achange in orientation within lattice cells which allows for an adaption of directional dependent elastic properties of TPMS lattice structures during design and optimization. To ultimately align the anisotropy of microstructures to stress direction, it is critical to release the rotational freedom of the unit lattice cell parameterization, in a manner similar to that done by Bendsoe and Kikuchi [27] in their homogenization method.

The distribution of lattice orientation in the design domain would allow for an adaption of each unit cell anisotropy to the local stress direction, rather than controlling the anisotropy by assembly two different base units to form a new representative unit as in [28].However, there some limitations on the spatial variant lattice construction. Theorientation can infill volume fraction distribution field should be continuous, the significant local distortion will result if the difference between neighboring cell angles is greater than 45°. For large angle difference, the standard triply symmetric TPMS structure will no longer exist in the distorted unit cell as the adaption of orientation difference. In this case, the neck area is elongated to offset the lattice distortion.

In this case, the neck area is elongated to offset the lattice distortion. Furthermore, the construction algorithm will lead to the lattice shift outside of the unit cell boundary to keep the required smooth transition due to the varying grating phase.The relationship between the variant volume fraction of the TPMS and iso-surfaceparameter t is not a linear one. Typically, the relative density of the unit lattice is controlled by changing the value of t to change the enclosed volume of the 3D surface or offsetting t to produce surface-based on lattice. Due to the introduction of infill fraction that is caused by the threshold method in the current algorithm, the relationship between t and volume fraction no longer exists. In this case, the relative density can be extracted from the summation of binary pixel by comparing pixel values with the threshold value.

In this case, the relative density can be extracted from the summation of binary pixel by comparing pixel values with the threshold value. A more accurate volume faction is obtained by recalculating the summation of the volumes of tetrahedron after meshing the enclose patch surface.Compared to the laminate TPMS lattice structure produced by recursive skewtransformation, the current lattice structure shifts out of the unit cell after spatial transformation. Due to the continuous orientation accumulation in one direction which is like the wave accumulation in propagation, the rotation of cells in Y direction leads to the skew shift in the X direction. If the orientation angle difference is small, this defect will disappear. However, the morphological variation generated by the spatially varying algorithm keeps a smooth transition while varying the lamina orientation, but the hybrid design with skew transformation does not.

Besides, if the relative volume fraction is lower, the connectivity between each layer is destroyed as shown in Fig. 4 (a).Since the spatial variant lattice needs to decompose to the planar grating vector in order to incorporate the varying orientation and solve for the grating phase, the presented research is limited to a primitive network or scaffold based TPMS lattice structure and is not applicable to surface-based lattices. There are two approaches to extend the current algorithm to other types of scaffold-based lattices, such as Gyroid, IWP and Diamond. From the generalized method, the1797orientation can be introduced to the truncated and decomposed Fourier terms to construct the mapping equation, which is not very computation efficient.

From the generalized method, the1797orientation can be introduced to the truncated and decomposed Fourier terms to construct the mapping equation, which is not very computation efficient. The more elegant approach takes less computation resources, since the unit TPMS lattice can be generated by the basic reciprocal vector as mentioned before [18], where the number of reciprocal vectors is far less than the truncated Fourier terms. The rotation matrices can be applied on the reciprocal vectors directly to construct the mapping equation and later solved by the finite difference method. With the solution of mapping function, the spatially varying TPMS lattices can be reconstructed by the generalized exponential expression.In addition, the present algorithm cannot distribute the TPMS lattice type simultaneouslyaccording to user-defined boundaries. Hence, two different methods were proposed to design heterogeneous TPMS lattice structures with varying orientation distribution.

Hence, two different methods were proposed to design heterogeneous TPMS lattice structures with varying orientation distribution. Unfortunately, the introduction of varying orientation within the lattices may lead to uncertain defects such as the sharp corners or unconnected ligaments when it comes to the heterogenous lattice structures.The transformed lattice under varying orientation needs to be combined with the cubicsymmetric TPMS lattice to produce the heterogenous lattice structures, where the defects between different types of lattices is unavoidable. The IWP type lattice is more compatible with Primitive type (similar with Body-centered Cubic) than Gyroid shown as Fig. 5 and Fig. 6, due to the BCC lattice (Primitive type) with varying orientation being closer to the geometric configuration of Face-Centered Cubic (FCC).

5 and Fig. 6, due to the BCC lattice (Primitive type) with varying orientation being closer to the geometric configuration of Face-Centered Cubic (FCC). In addition, the transition area produced by the distance field function shows a better geometry connectivity, however, it is necessary to evaluate its construction quality by finite element analysis in the future research.5. ConclusionIn summary, a novel method is proposed to design spatially varying TPMS lattice structures with varying infill volume fraction and varying spatial orientation. An iterative central finite difference method was applied to solve the partial differential equation governing the relationship between varying orientation and position. The current method can be utilized to optimized directional dependent TPMS lattice structure to align principal stress with the lattice geometry configuration within the design domain.

The current method can be utilized to optimized directional dependent TPMS lattice structure to align principal stress with the lattice geometry configuration within the design domain. In addition, the heterogeneous lattice with varying orientation is explored by GRBF and the distance field function where the later one is shown to produce better structure transitions between different lattice types. By introducing the lattice varying volume fraction, unit cell variant orientation and lattice type distribution, more freedom is offered to the design domain of TPMS lattice structures.1798Reference[1][2][3][4][5][6][7][8][9]Alabort, Enrique, Daniel Barba, and Roger C. Reed. "Design of metallic bone by additive manufacturing." Scripta Materialia 164 (2019): 110-114. Arabnejad, Sajad, et al. "Fully porous 3D printed titanium femoral stem to reduce stress‐ shielding following total hip arthroplasty." Journal of Orthopaedic Research 35.8 (2017): 1774-1783.

Multi-morphology transition hybridization CAD design of minimal surface porous structures for use in tissue engineering. Comput. Aided Des. 56, 11–21[10] Yoo, Dongjin. "Heterogeneous minimal surface porous scaffold design using the distancefield and radial basis functions." Medical engineering & physics 34.5 (2012): 625-639.[11] Afshar, M., et al. "Additive manufacturing and mechanical characterization of graded[12]porosity scaffolds designed based on triply periodic minimal surface architectures." Journal of the mechanical behavior of biomedical materials 62 (2016): 481-494. Cheikho, K., Cédric Laurent, and J. F. Ganghoffer. "An advanced method to design graded cylindrical scaffolds with versatile effective cross-sectional mechanical properties." Journal of the mechanical behavior of biomedical materials 125 (2022): 104887.[13] Ma, Songhua, et al.

sign of Stereolithography Trees for Use in the Investment Casting of Stereolithography Patterns Christopher S. McDowell*, Mark C. Boomer Johnson & Johnson Professional, Inc., 325 Paramount Dr., Raynham, MA 02767Investment Casting Stereolithography PatternsStereolithography (SL) has become a useful tool in rapid prototyping of metal parts due to advances in resins and buildstyles for producing investment casting patterns. The production of the patterns, however, has been only one obstacle to the successful application of SL to investment casting. Difficulties in pattern removal from investment casting shells has been a major concern for foundries handling SL patterns, limiting even the most experienced foundries to 90% In addition to shell cracking, improper and inadequate gating of SL patterns can lead to yields. casting defects or failure to produce a satisfactory part.

casting defects or failure to produce a satisfactory part. The costs of SL pattern generation, shell production, and raw metal stock, and the tight time requirements of rapid prototyping demand that higher foundry yields be achieved.The investment casting process begins with the generation of a pattern. The pattern reflects the desired part geometry with an allowance for metal shrinkage. One or more patterns are attached to a structure which will provide a pathway for the molten metal to reach the patterns. This structure is a known as a tree and consists of feed tubes that carry molten metal to one or more gates. Gates connect the feed tubes to the patterns. Once the SL patterns are affixed to the tree, a ceramic shell is created around the tree and patterns. The shell is formed by repeatedly dipping the tree/pattern assembly into a slurry of ceramic and allowing each coat to dry before repeating the process.

The shell is formed by repeatedly dipping the tree/pattern assembly into a slurry of ceramic and allowing each coat to dry before repeating the process. Sand is frequently poured over the ceramic before it dries to add strength to the shell. Once the shell is complete, the pattern and tree are removed from the shell by melting or burning. The shell may then be filled with molten metal, which solidifies in the cavities left behind by the patterns. Finally, the shell is broken away and the metal parts are cut from the tree.3D-Systems' (Valencia, CA) development of the QuickCast™ (QC) buildstyle has made the rapid generation of accurate investment casting patterns possible. The use of traditional wax trees in investment casting SL patterns, however, has limited the scope of SL in investment casting to pattern generation. This paper focuses on the use of SL to create trees for investment casting, and explores the possibility of using SL trees instead of wax trees.

This paper focuses on the use of SL to create trees for investment casting, and explores the possibility of using SL trees instead of wax trees. Casting SL patterns and trees may be effective in improving foundry yields, reducing the amount of metal required to create a casting, and even improving the quality of the resulting castings.313Applications of SL treesPattem geometry has a significant impact on the degree of difficulty in investment casting SL patterns. l'hin and narrow features on SL patterns act as stress risers during pattern removal and may lead to shell failure. Thick sections and large geometries, while presenting less of a problem in t:ermlrS of pattern removal, are difficult to adequately feed and may require strategic placement ofieed tubes and gates.

Thick sections and large geometries, while presenting less of a problem in t:ermlrS of pattern removal, are difficult to adequately feed and may require strategic placement ofieed tubes and gates. Development of SL trees for investment casting may address both of these aspects of investment casting SL patterns by making tree geometries flexible and by eliminating the use of autoclaving as part of the pattern removal process. Parametric Technology'.s Pro/Engineer® (Waltham, MA) solid modeling software was used in conjunction with 3D-Systems' SLA 250/40™ and Maestro™ software to create the trees used in this work.Preventing Slid Cracking During Pattern RemovalPattern. removal represents the most significant problem arising from the use of QC patterns. Autoclaving is the traditional investment casting wax removal process. However, QC patterns exhibit extreme swelling upon exposure to steam during autoclaving, due primarily to the absorption of water.

However, QC patterns exhibit extreme swelling upon exposure to steam during autoclaving, due primarily to the absorption of water. Part geometries such as thin walls, small holes, and sharp corners cannot readily collapse inward as this swelling occurs, resulting in shell cracking around the patterns. To avoid this problem, most foundries bypass the autoclaving step and subject shells containing SL patterns to direct burnout of the patterns in a flash-fire oven, typically the mold-preheat furnace.While work continues on resins that are less sensitive to moisture, flash-firing remains the most promising route towards removal of patterns with difficult geometries. The use of the flash fire technique, however, merely relocates the potential for shell failure when wax trees are used.

The use of the flash fire technique, however, merely relocates the potential for shell failure when wax trees are used. Although the QC patterns generally have a greater coefficient of thermal expansion than the wax!, their relatively open internal structure allows the pattern to fail and collapse inward during burnout, before sufficient stress is generated to crack the shell. The expansion of the solid wax tree, however, can build up tremendous pressure along the tree and lead to shell failure.

The expansion of the solid wax tree, however, can build up tremendous pressure along the tree and lead to shell failure. It is important to note that some SL pattern geometries necessitate the use of a thick shell, in which case the use of a wax tree may not be precluded.Changing shell composition to improve shell strength2 may allow the processing of SL patterns without a substantial for foundries whose processing of SL parts is only a small fraction of overall part production, or as in regulated industries where changes to shell composition require lengthy validation procedures, changing shell compositions to handle SL patterns is unrealistic. Increasing shell strength by increasing shell thickness is possible, but at significant cost due to several factors. First, an increase in shell thickness typically requires patterns to be spaced farther apart, decreasing the number of patterns that can be applied to the tree.

First, an increase in shell thickness typically requires patterns to be spaced farther apart, decreasing the number of patterns that can be applied to the tree. Second, the additional process time required to augment the shell thickness adds to the total time required to produce the shell. As time is an essential element in RP processing, such an increase may be unacceptable. Finally, the increased usage of shell materials adds cost to the tree, as does the additional labor required for processing.increase in shellthickness. However,314When pattern geometry requires avoiding autoc1aving, it is possible to create an SL tree to which SL patterns may be attached prior to burnout and which provides better survivability of the shell.

A properly designed SL tree can also improve the circulation of combustion gasses, a critical factor in the successful burnout of SL patterns3.Key to the successful use of the SL tree for reducing shell cracking is the incorporation of features designed to facilitate the collapse of the tree during the initial moments of burnout. The primary principle of SL tree design is the use of the ACES™, or solid, buildstyle, and the generation of a thin-walled, hollow geometry. This geometry can reproduce the desired shell volume while using less SL resin than an identical part made in the QC buildstyle. Additionally, thin walled part collapses at a lower stress during burnout and requires less oxygen to completely combust, freeing up the combustion of the patterns. Since a good surface finish on the tree is not critical, the tree can be built 0.006" thick layers. QC patterns typically built using are 0.004" to improve surface finish.

Since a good surface finish on the tree is not critical, the tree can be built 0.006" thick layers. QC patterns typically built using are 0.004" to improve surface finish. As a result of the larger layer thickness, the time to build is an substantially less than the time it would take to build the same tree as a QC pattern.theoxygenforthicklayersACEStreeFigure 1 - Wax tree used for investment casting of production and prototype parts. Scissors are for scale.Figure 1 shows a wax tree currently used to cast orthopedic implants in production. The tree is injection molded with a threaded insert in the mold that allows a threaded rod to be screwed into the tree for manipulation during dipping. This tree was used in conjunction with QC patterns to cast prototype parts in ASTM F-75, a Cobalt-Chromium-Molybdenum alloy. The components had thin sections and sharp edges, and it was difficult to drain excess resin from the patterns during processing.

The components had thin sections and sharp edges, and it was difficult to drain excess resin from the patterns during processing. Autoclaving the mold to remove the SL patterns and wax tree proved disastrous; the ceramic mold was completely destroyed by the expansion of the SL patterns. Extensive shell cracking was also observed after flash-firing the mold to remove the patterns and wax.315Shell failure caused by flash-firing was observed along the length of the tree, and was to stresses generated by the thermal expansion of the wax during the early stages of burnout. An attempt to remove the wax by slow heating (with the intent of subsequently burning out the SL patterns) a convection oven also led to shell failure. To alleviate the cracking problem, an SL tree was created that mimicked external the geometry of the wax tree.attributedinA cross section of the SL tree is shown in several Figure 2, whereFigure 2 - Section through an SL tree.

The SL tree reproduces the external geometry of the wax tree shown in Figure 1.important features of the tree are revealed. Perhaps the most notable feature of the tree is its thin walls (A). At 0.025" thick, the walls of the tree were strong enough to hold QC patterns general and handling during shelling, but thin and weak enough to break readily as stresses developed firing. All corners and edges were filleted and rounded to stress reduce concentrations the ceramic shell.surviveduringinFigure 3 - Section of SLA tree showing support structure for threaded insert.notable feature of the tree is the joint connecting the two portions of the tree. The tree was too large toAnother(B)316be built in a single piece on the SLA 250, so it was split in ProlEngineer prior to building. The joint needed to be strong enough to hold the weight of the patterns on the lower half of the tree.

The joint needed to be strong enough to hold the weight of the patterns on the lower half of the tree. The joint, a "bird's beak" type of connection, was sufficiently strong to hold the two parts together. The benefit of this type of joint is that it collapses into a narrowing "v" shape as circumferential stresses arise in the tree during firing. A simple flange-type of joint would be more structurally resistant to failure during burnout, transmitting stress to the shell and potentially leading to shell failure. Figure 3 shows a close-up of the top portion of the SL tree of Figure 2. A reinforcing structure was added to the tree to accommodate the threaded insert (shown in place) which was present in the original wax tree. The insert was placed into the tree and rotated 45 degrees to lock it into position under the reinforcement. The structure was made independent of the wall of the tree, and was strong enough to support the weight of the shelled tree.

The structure was made independent of the wall of the tree, and was strong enough to support the weight of the shelled tree. The strength of the insert reinforcement was not transmitted to the outer wall of the SL tree, preventing the transfer of additional stress to the ceramic shell during firing.Figure 4 shows cross sections through two flash-fired ceramic molds made from the tree geometry shown in Figure 1. To create a ceramic shell strong enough to withstand flash firing of a wax tree required 11 slurry/sand layers. A section through the fired shell is shown in Figure 4a. The shell in Figure 4b was created over the SL tree shown in Figure 2. The shell over the SL tree survived firing even though only 6 slurry/sand layers were applied.

The shell in Figure 4b was created over the SL tree shown in Figure 2. The shell over the SL tree survived firing even though only 6 slurry/sand layers were applied. Producing the ceramic mold by using an SL tree saved 2 days in the shell production area and reduced the weight of materials (slurry & sand) by 66%.Figure 4 - a) Fired and sectioned investment casting shell showing thickness of ceramic mold created over a wax tree. b) Fired and sectioned shell showing thickness of ceramic mold created over a SL tree. The 1.D. of the shells was the same.Optimizing the Investment Casting of SL PatternsDue to the large variety of pattern geometries encountered in prototype development, the gating and feeding of SL patterns remains, largely, an art. Foundries rely on experience for ensuring adequate fill of the mold as well as control over the actual filling and solidification of the317metal.

Foundries rely on experience for ensuring adequate fill of the mold as well as control over the actual filling and solidification of the317metal. The usual strategy for feeding these parts is to apply excess gating, accepting less than optimal metal flow and extra finishing work on the castings. As an alternative, development of an investment casting tree might be performed in conjunction with the original solid model for the part. This would promote visualization of the entire casting process early on, perhaps eliciting design changes that improve the castability of the part.When prototype parts need to be cast from expensive alloys, the elimination of metal waste becomes an important criterion for designing the tree.

Using feed tubes and gates with circular cross sections minimizes heat loss in the shell while the metal is being poured, allowing the use of smaller diameter feed tubes and gates.Since part complexity is not generally an issue in SL fabrication, complex features may be added to improve the metal flow characteristics of the mold. For instance, the outer surface of the tree might be grooved slightly to promote the laminarity of metal flow through the ceramic mold upon casting. Reducing turbulent metal flow through molds can reduce the chance that small ceramic fragments in the mold become trapped in the casting. Smooth gate-part transitions and bends in the tree can also help to promote laminarity.When part geometries contain thick cross sections, metal shrinkage can cause voids to appear in the castings.

Smooth gate-part transitions and bends in the tree can also help to promote laminarity.When part geometries contain thick cross sections, metal shrinkage can cause voids to appear in the castings. To avoid shrinkage cavities, a reservoir of molten metal called a riser is created by forming a volume in the shell that solidifies after the part, allowing additional molten metal from the riser to fill the part cavity as solidification takes place. To keep the metal in risers molten, they must often be very large. Maximum heat retention is accomplished when the surface area of the mold is small in relation to the mass of metal in the riser. By using SL, hollow cylindrical or semi-spherical risers (geometries with low surface area/volume ratios) could be created and added to the tree to prevent shrinkage cavities.Provisions must generally be made for handling the tree during the shelling process.

In the example above, a special feature was made for allowing a threaded insert to be placed into the SL tree. This insert must be removed prior to burning out the shell. As an alternative to a separate, threaded insert, screw threads could be modeled directly into the SL tree.Determining the Cost of SL TreesAs the production of SL parts can be expensive, minimizing the use of SL resources, including resin, build time, and post processing requirements is critical in making the use of SL trees feasible. Since surface finish and dimensional accuracy of SL trees is ,relatively unimportant, trees should generally be positioned to minimize build height. In the example described earlier, the lower, long portion of the SL tree was built lying on its side. Supports were generated inside the part, however, they did not contribute significantly to the strength of the part and were left in place (reducing the amount of post processing required).

Supports were generated inside the part, however, they did not contribute significantly to the strength of the part and were left in place (reducing the amount of post processing required). The total part height was reduced to about 2 inches, and though the tree was slightly out-of-round (the effect of cure depth and overcure characteristic of the SL process) there were no adverse effects on the performance of the318tree. The tree consumed only 60g of resin (about $4 worth ofresin).*The small cross-sectional areas of these SL trees require minimal scan time. SL trees do not require any special build parameters, and can be placed into a build when there is extra room in the SL apparatus, to be built along with other ~ACES style parts. About 3 hours were required to model the SL tree from the above example in ProlEngineer.SL trees may also be used for proofing out tree geometries prior to producing production tooling for injecting wax trees.

Optimization of tree geometry in the production of large numbers of castings can have tremendous impact on production costs. The flexibility offered by solid modelling and SL technology allows timely experimentation with a number of parameters of tree geometry that ultimately affect the raw metal requirements and quality ofthe castings.SL trees clearly provide benefits over wax trees in reducing shell cracking when SL patterns with difficult geometries are to be cast. SL trees may also be employed to reduce shell material and process time, improve foundry yields, reduce raw metal stock requirements, and quality of the casting, providing ample economic return on the investment of improve the producing the SL tree.ConclusionsThe demanding pace of rapid prototyping and the high cost of SL patterns require higher foundry yields than are currently achieved.

The use of SL trees in the investment casting of SL patterns has the potential to improve foundry yields significantly, through improved gating of parts and a reduction in shell failures during pattern removal. The SL ACES buildstyle in conjunction with solid modeling software packages can be used to create thin-walled tree geometries possessing an array of features that can significantly reduce shell stresses developed during flash-fire burnout and improve the flow of metal into the mold. The time savings associated with improved foundry yields and elimination of the autoc1aving step can be one of the most significant advantages of the use of SL trees. While the initial cost of designing and producing an SL tree may be higher than that to produce a wax tree, the potential to ensure shell survival during pattern removal, conserve expensive alloys, and produce high quality castings may justifY the added expense.References1. K.

K. Losse, ""Investigation of Optimum Process Parameters for Burnout of SLA Parts for Manufacturing of Metallic Parts Through Investment Casting," Proceedings, Prototype Research and Manufacturing '96, 5MB, Dearborn, MI April 23-25, 1996.2. M. Guerra, W.D. Roberts, and A. T. Bozzo, "Factors Affecting Shell Strength and the Effect of Dry Time on Shell Strength," 22nd EICF Conference, Paris, 22-25 April, 1992. 3. C. Atwood et aI, "Producing Castings from Rapid Prototyping Patterns," presented at Prototype Research and Manufacturing '96, 8MB, Dearborn, MI April 22, 1996.Due to the wide variation in pricing of SL parts between service bureaus and in-house facilities, no attempt will be made here to estimate an overall cost for producing the tree.319320.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesign Optimisation of a Thermoplastic SplintAngus Fitzpatrick, Mazher Mohammed, Paul Collins, Ian GibsonDeakin University, School of Engineering, 75 Pigdons Rd, Waurn Ponds, Geelong, Vic 3216, AustraliaKey words: Patient Specific, Design, 3D printing, Orthosis, OptimisationAbstractFollowing partial hand amputation, a post-surgery orthosis is required to hold the remaining ligaments and appendages of the patient in a fixed position to aid recovery. This type of orthosis is traditionally handmade and fabricated using laborious and qualitative techniques, which would benefit from the enhancements offered by modern 3D technologies.

This type of orthosis is traditionally handmade and fabricated using laborious and qualitative techniques, which would benefit from the enhancements offered by modern 3D technologies. This study investigated the use of optical laser scanning, Computer Aided Design (CAD) and Material Extrusion (ME) additive manufacturing to manufacture a polymeric splint for use in post-surgical hand amputation. To examine the efficacy of our techniques, we take an existing splint from a patient and use this as the template data for production. We found this approach to be a highly effective means of rapidly reproducing the major surface contours of the orthosis while allowing for the introduction of advanced design features for improved aesthetics, alongside reduced material consumption. Our demonstrated techniques resulted in a more lightweight and lower cost device, while the design and manufacturing elements afford greater flexibility for orthosis customisation.

Our demonstrated techniques resulted in a more lightweight and lower cost device, while the design and manufacturing elements afford greater flexibility for orthosis customisation. Ultimately, this approach provides an optimized and complete methodology for orthosis production.IntroductionThe use of orthotics devices post-surgery is an important part of the overall patient recovery process, particularly with respect to cases involving the upper limbs, such as hand/arm amputations (Carl R. Chudnofsky 2010) and shoulder joint replacements (Browner et al. 2014). The orthosis allows for vital load bearing support of the patient’s residual appendages and whilst their remaining ligaments/musculature recover (Cross 2014).

2014). The orthosis allows for vital load bearing support of the patient’s residual appendages and whilst their remaining ligaments/musculature recover (Cross 2014). Traditionally, such orthotic devices are custom made, using thermoforming plastics, conforming to the natural anatomy of the patient, providing a configuration to stabilise the residual appendages in a desired orientation for directed recovery (MacDonald, Chinchalkar & Pipicelli 2016).Modern 3D technologies are now allowing for unprecedented accuracy and flexibility for product development, with benefits to be seen in terms of 3D data acquisition (Treleaven & Wells 2007), how design iterations can be explored and developed and the final manufacturing of the prototype and functional devices (Bagaria 2015; Mohammed et al. 2016).

2016). Given the overwhelming potential of this technology as a transformative product development paradigm, we have seen adoption across multiple fields ranging from sports technology, prototyping of commercial products, and with a considerably large uptake for medical applications (Mohammed, Fitzpatrick & Gibson 2017; Mohammed et al. 2017). More recently, the use of such technology for orthotic development and production is a highly emerging area (de Jesus Faria 2017; Fitzpatrick et al. 2017; Kelly 2015; Lasane 2013; Paterson et al. 2015; Summit 2007), but also is a largely underexploited area of interest. We2409predict with the rise in the net global, and in particular aging, populations, such developments would become of growing interest and necessity in the coming decades.In this study, we aim to develop a streamlined methodology for next generation orthotics production which leverages the advantages of emerging 3D technologies.

As a case study, we examine an existing orthotic device made for a patient who has undergone a partial hand amputation, assess the merits and limitations both in terms of efficacy, ergonomics and aesthetics, and apply a combination of optical scanning, CAD and additive manufacturing to redevelop the part to address these limitations. We discovered optical scanning allows for rapid virtual reproduction of the orthosis (1-2minutes). The use of CAD can allow for the incorporation of intricate open structures alongside part thickness tolerance considerations; something not readily achievable using traditional fabrication techniques. Finally, additive manufacturing using ABS Material Extrusion (ME) printing readily reproduces the intricacies of the CAD design, producing a robust and superior final orthosis.

Finally, additive manufacturing using ABS Material Extrusion (ME) printing readily reproduces the intricacies of the CAD design, producing a robust and superior final orthosis. We believe our findings would provide guidance to orthotists to streamline partial amputation orthotics and potentially provide a range of alternative orthotics for both the upper and lower limbs.Original orthotic assessmentIn the present case study, we examined an orthotic device made for a patient who had undergone a partial amputation of their hand (thumb and index finger). The devised splint was required post-surgery to help support and fix the mobility of the ligaments, nerves, and tendons to aid in the recovery process. Such devices are normally made using a thermoformable plastic, made of cured polyester resin, which has been the industry standard for several decades (fig 1).

Such devices are normally made using a thermoformable plastic, made of cured polyester resin, which has been the industry standard for several decades (fig 1). The plastic is available in a range of thicknesses (1.6-5.0mm), and can have small holes regularly spaced within the bulk of the material, with the intension of providing porosity for moisture release.It is typically found however that such orthotic devices made from this material are generally quite uncomfortable for patients, due to both weight and excessive moisture build- up (even when using the perforated version), which causes issues such as chaffing. In worst- case scenarios, excessive moisture build-up and chaffing can result in ulceration (Gregory P. Guyton 2005). Additionally, the aesthetics of the device generally makes patients feel highly self-conscious of their condition, resulting in issues with respect to wider societal inclusion.

Guyton 2005). Additionally, the aesthetics of the device generally makes patients feel highly self-conscious of their condition, resulting in issues with respect to wider societal inclusion. We therefore identified that modifications could be made to the design to reduce the weight (in this case approximately 110g) to reduce strain on already compromised musculature, while introducing design elements, such as open structures, to both minimise moisture retention and improve the overall aesthetics.Fig 1. A diagram illustrating the typical phases of traditional upper limb orthotic production2410Methodology Optical Scanning and DigitisationThe current thermoplastic splint was virtualised using a Kreon Solano 3D laser scanner (Kreon, Belgium), attached to a 7-axis rotatable arm for ease of scanner translation. The scanner uses a red laser line, with a wavelength of 610-630nm (George et al. 2016), and a triangulation method of measurement.

The scanner uses a red laser line, with a wavelength of 610-630nm (George et al. 2016), and a triangulation method of measurement. To perform a scan, the laser line is translated over the object at a fixed distance to build up the point cloud data profile for a respective surface. Once an exposed surface has been recorded, the orthosis was re-orientated to scan additional surfaces until the entire profile was obtained.The scanner creates a point-cloud data file within its proprietary software, which requires post-processing to remove any outlying data points and any unwanted additional scan data from the surrounding environment. Once post-processing is complete the point-cloud forms the basis of the final 3D model. After cleaning, the dataset was solidified into an STL (Stereolithography) format. The STL model was measured virtually and compared to measurements taken from the physical model.

After cleaning, the dataset was solidified into an STL (Stereolithography) format. The STL model was measured virtually and compared to measurements taken from the physical model. The exported STL file was then taken into 3- Matic STL (Materialise, Belgium) for further error correction to ensure model integrity for 3D printing. Following this procedure, the virtual model was again checked for dimensional accuracy.Orthotic DesignThe exported model, once taken into 3-Matic STL, is in the same orientation as the scan was taken. Designs were developed using the original orthosis topography as the basis. At this stage, we use the data to introduce a parametric tile pattern (Figure 3.2) into the bulk of the digital structure. The pattern is scaled according strength and aesthetic requirements and it must be ensured that there is an effective boundary. To create this Surface map, a surface contour must be defined. This Surface is used to create a UV map (Policarpo et al. 2005).

To create this Surface map, a surface contour must be defined. This Surface is used to create a UV map (Policarpo et al. 2005). Using the UV mapping process, the chosen design can be applied to the surface in 3-Matic STL (Fitzpatrick et al. 2017). Using the applied map, the design can be subtracted from the blank model using the texture function, removing any unwanted sections. As the applied map is an image file, the borders of the pattern edges are jagged, this is due to the subtraction process as well as the tessellation of the model prior to this operation. This is fixed using the surface smooth function in the software.Next the model is again checked for accuracy to ensure the design process has not moved any of the external faces which may affect comfort. Following this, FEA (Finite Element Analysis) is performed on the design to ensure mechanical integrity for the intended application.

Following this, FEA (Finite Element Analysis) is performed on the design to ensure mechanical integrity for the intended application. The variables that could change within the design are the splint thickness and the spar thickness. This could be an iterative process that would involve a dialog between the designer and the client. This is then verified for 3D printing and exported to the slicing software.ManufacturingWe opted for polymer-based ME as the preferred additive manufacturing method, which allowed for accurate reproduction of the developed design, whilst also being suitably low cost to serve as a method for mainstream orthosis production. The final design was printed2411using ABS plastic, due to its low cost and suitable mechanical and biocompatibility properties, and with the orthotic device orientated upright in the Z build direction. Initially, preliminary design iterations were printed using a Zortrax M200 ME printer (Zortrax, Poland).

Initially, preliminary design iterations were printed using a Zortrax M200 ME printer (Zortrax, Poland). However, due to build volume constraints, the optimum orientation could not be reached, in addition to printed models being relatively weak. It did however allow a test fit model to be created at a much lower cost than the final splint. To overcome the mechanical and size constraints a Fortus 450M (Stratasys, USA) FDM printer was utilised to produce the final orthosis.Results VirtualisationThe time taken to 3D scan the splint accurately was around 45 minutes.This was because the scanning arm was unwieldy to use. A turn table would have made it easier, but it wouldn’t have worked with the current scanner as it uses the arm to register its position rather than the surface topology. The aim was to get the splint scanned in one pass, rather than stitching multiple scans together.

The aim was to get the splint scanned in one pass, rather than stitching multiple scans together. Although more time-consuming, this made the result more accurate (see fig 2).The accuracy of the virtual model produced was +/- 0.1mm using a scanner resolutionof around 16µm, measured within 3-Matic.(a)(b)Fig 2. The initial scan of the splint and the resultant cleaned modelThe process of virtualisation was made easier as it was scanning an inanimate object.It would be made inherently more difficult if the scan was of a human subject as any movement would be shown in the scan.Design process The integration of end user input is an important aspect in customised designs, such as what we hope to achieve with the developed splint design. This is a trend that has been increasing within other fields of engineering (Collins 2016). However, this is a concept that has had a hesitant uptake in the medical devices industry (Money et al.

This is a trend that has been increasing within other fields of engineering (Collins 2016). However, this is a concept that has had a hesitant uptake in the medical devices industry (Money et al. 2011), but there are indications that this is beginning to change (de Jesus Faria 2017). Having control over the design of the treatment device has been shown to increase the quality of life (Fayers & Machin 2013).2412(a)(b)Fig 3. Design variants chosen by the end user.To examine facets of user based design, we selected two arbitrary, yet aesthetically pleasing designs (Fig 3) from the perspective of the researchers in this study, and assessed for viability for the splint. From preliminary design evaluations, the pattern in Fig 3(b) was chosen for the final design as it provided a greater strength throughout the splint. Additionally, this design was advantageous as manufacturing required less support material and could be achieved comparatively more quickly.Fig 4.

Additionally, this design was advantageous as manufacturing required less support material and could be achieved comparatively more quickly.Fig 4. Surface contour defined for the virtual model.The mapping process first requires the definition of the surface, leaving a borderaround the external contours of the virtual splint, shown in fig 4. The reasoning for this is to allow enough room for an adequate border around the splint to ensure its strength and stability.The pattern is then applied to the UV map. This is then resized, positioned and rotateduntil it is in the correct orientation. Shown in fig 5, the UV map is shown alongside the wrapped model. This allows the accurate positioning of the pattern and to ensure the borders will have adequate support. The thinking behind the border was to minimise the sectioned holes, as they would be filled in the next step to increase the size of the splint border and to integrate the design seamlessly.2413Fig 5.

The wrapped model shown alongside the UV map.Following the texture application, the model is modified using the 2D to 3D texture function in 3-Matic STL. The function gives the user the options of offsetting the surfaces positively or negatively in the direction normal to the surface. Using the external surface of the splint for the texture mapping, the white sections were offset through the splint, while the blue surfaces remained in a fixed position, as shown in Figure 6(a). At this stage, the model was error checked again, to ensure no digital anomalies (inverted normal, overlapping triangles, etc). The resulting splint model is shown in Figure 6(b).(a)(b)Fig 6. Splint overlayed with a textured model and the final modelThe accuracy of the model following the design application stays the same as the scanned model. This is because borders and datum points were used to ensure any design variations wouldn’t affect the applicability of the model.

This is because borders and datum points were used to ensure any design variations wouldn’t affect the applicability of the model. Any changes that were made to the splint were completed on a duplication of the original scan. This was to assess the accuracy of the model through the design period.This is completed in both 3-Matic STL as well as Magics (Materialise, Belgium),allowing for a smaller file size and the slicing software to quickly process the model.Additive Manufacturing allows more complex designs to be manufactured. Using MEto print the model, the design needs to be as self-supporting as possible to minimise the material used in manufacture.2414Following exporting the STL files, renders were done in Showcase (Autodesk, USA) for colour and design approval, shown in fig 7. Additive manufacturing allows the production of multiple pieces in a variety of colours, giving the end user the ability to pair the splint to their outfit for example.(a)(b) Fig 7.

Additive manufacturing allows the production of multiple pieces in a variety of colours, giving the end user the ability to pair the splint to their outfit for example.(a)(b) Fig 7. Resultant models in colour variants(c)The current splint design was constrained by the aesthetic and ergonomic properties of the thermoplastic splint, however performed relatively satisfactory from a biomechanical stand point. For this reason, we aimed to retain these aspects such that the overall longitudinal rigidity, rotational stiffness and latitudinal compressive strength were equivalent in the developed splint. As the pattern in Fig 3(b) would result in a more open, porous structure in the bulk of the splint, we needed to ensure that all compressive or shear forces would be even distributed throughout the design.

As our primary intension in this study was to demonstrate the design and product realisation facets of the splint, we hope to address this in future studies.ManufacturingThe Splint took 23 hours to print, with 58g of ABS. The print was completed with the splint lying flat of the print bed. This was due to height restrictions in the Fortus. Because of the orientation, the splint required support. The weight of support used was 50g contributing to a slower build time.The initial model printed on the Zortrax, and the final part, completed on the Fortus,(shown in fig 8.), acted very differently under load, with the Zortax model fracturing relatively early. This could have been due to the orientation that it was printed at. More work will be completed to characterise each of the designs and the mechanical properties that are affected by design changes.2415Fig 8.

More work will be completed to characterise each of the designs and the mechanical properties that are affected by design changes.2415Fig 8. Result after manufactureDiscussion and future workComparing the thermoplastic and 3D printed splints, some of the notable differencesbetween these approaches are;The reduction in weight, with a possible 60% reduction - The breathability of the AM splint should be considerably better in comparison to the thermoplastic, due to the open pore design.The comfort of the splint, due to the reduction in weight and sweat build up, should be considerably better with the AM splint.Furthermore, indications are that the splint should chaff less as the design doesn’t appear to move in relation to the body as much.

The aesthetic appeal of the splint can be adjusted to suit the user, with even the possibility of having a number of variants to suit aesthetic requirementsFuture work planned will take a more quantitative perspective, looking at the engineering properties of the design variants, material properties, stress, fracture load, etc. Being able to characterise the splint properties by varying the design constraints should provide a better quality of life for the end user, as well as make it easier to manufacture the splints. A variety of designs can be assessed, allowing the end user to have a wider range of patterns that can be easily designed and manufactured, while ensuring the correct care is provided.The major problem to be solved is reducing the overall time to create the orthosisfrom start to finish. A more effective scanning method is being developed that will scan the upper limbs in real time.

A more effective scanning method is being developed that will scan the upper limbs in real time. We are also working on optimising the design process to streamline the scan data manipulation whilst incorporating the aesthetic design components and maintaining mechanical integrity. Finally, we have chosen a Delta configuration ME system (Anzalone et al. 2015) which will allow the construction of larger orthoses to suit the upper limb with minimal supports. All of these approaches will considerably speed up the design and build process.The materials that have been used in this study, will be assessed and compared in a future study. The study will look at the variation on materials with a standard design. These results will be compared to other materials that are used in current splinting practices.2416Conclusion Concluding this study, it was found that the current thermoplastic splint, whilsteffective, it does not seem to be a perfect solution.

The use of 3D printed splints can enable tailoring to suit the end-user’s preferences, potentially increasing the quality of life during the treatment process. This can aid in the end-user’s recovery as well as allowing the return to normality without the stresses of wearing something that they are not content with.ReferencesAnzalone, GC, Eujin Pei, D, Wijnen, B & Pearce, JM 2015, 'Multi-material additive and subtractive prosumer digital fabrication with a free and open-source convertible delta RepRap 3-D printer', Rapid Prototyping Journal, vol. 21, no. 5, pp. 506-19.Bagaria, V 2015, 'Technical Note: 3D Printing and Developing Patient Optimized Rehabilitation Tools (Port) - A Technological Leap', International Journal of Neurorehabilitation, vol. 02, no. 03, pp. 1085-9.Browner, BD, Jupiter, JB, Krettek, C & Anderson, PA 2014, Skeletal Trauma E-Book, Elsevier Health Sciences.Carl R.

ternational Solid Freeform Fabrication Symposium Aug 3-5, 2009, Austin TexasDesign Optimization of a Customized Dental Implant Manufactured via Electron Beam Melting®Gilbert Chahine1, Hosein Atharifar2, Pauline Smith3 and Radovan Kovacevic1 1: Research Center for Advanced Manufacturing (RCAM), Department of Mechanical Engineering, Southern Methodist University, Dallas Texas 7520 2: Department of Industry and Technology, Millersville University, Millersville Pennsylvania 17551 3: Army Research Laboratory, Aberdeen Proving Ground, MD 21005Finite Element Analysis (FEA) is a commonly used tool to evaluate biomechanics of traditional dental implants. Biomechanics help predict bone response and implant retention which strongly affects the longevity of the implant. The current research utilizes an analogues approach with FEA, to evaluate the biomechanics of a customized dental implant design built by Electron Beam Melting®, and to contribute towards the implant’s design optimization.

The analysis consists of three distinct simulation models. The first model is established in order to get an insight of the biomechanics produced by a biting force of 400 N on a second human molar in the mandible, its corresponding superposed mate and the surrounding biomaterial. In the second model, the lower jaw molar is replaced by a Ti-6Al-4V customized dental implant with a solid surface at the root. In the third model, the customized dental implant has a modified outer-layer at the root with adjustable elasticity.

In the third model, the customized dental implant has a modified outer-layer at the root with adjustable elasticity. By using a deterministic optimization technique in the FEA, an elasticity of the modified layer can be selected in a manner to minimize stress shielding from occurring.Introduction: In modern dental Implantology, dental implants are composed of three components; a screw-like root form inserted into a drilled and bored hole in the jawbone, an abutment which provides support for a dental prosthesis, and a screw which connects these components [1].Previous work discusses the new concept of a one-component customized dental implant which mimics the shape of the natural tooth and makes use of the socket produced by extraction [2]. The new concept potentially evades the need for drilling and boring used in traditional implantation, which consequently decreases mechanical and thermal trauma; hence, inducing faster healing time [2].

The literature discusses a case study, where a CT scan is obtained from a patient, and is used to design a customized dental implant. The design is manufactured using an EBM® A2 ARCAM® machine out of Ti-6Al-4V ELI.The biomechanics of dental implants are essential to bone response, implant longevity and stability. Consequently, it is important to compare the biomechanics produced by the use of customized dental implants to the biomechanics present in natural dentitions. Finite Element Analysis (FEA) is a cost- effective and insightful tool used in predicting mechanical and thermal behavior of traditional dental implants in their bio-environment. Such analysis enables researchers to obtain information regarding the nature of the biting load, and the induced stress and strain into the dental implant and the surrounding bone [3].

Such analysis enables researchers to obtain information regarding the nature of the biting load, and the induced stress and strain into the dental implant and the surrounding bone [3]. With the numerous benefits that FEA provides, similar analysis is used to evaluate the biomechanics of customized dental implants, and provide a gateway to optimize the mechanical features of their design.Electron Beam Melting® (EBM®): EBM® is a rapid manufacturing technique introduced and developed by ARCAM® AB. The technique uses a 4 KW electron beam gun to melt thin layers of metal powder under a controlled vacuum environment. The latest model; the EBM A2 is equipped with a larger and more consistent power supply when compared to its predecessors, as well as two interchangeable build chambers and more optimized software. The EBM A2 produces parts with excellent mechanical properties (comparable to wrought material) as well high resolution and part accuracy.

The EBM A2 produces parts with excellent mechanical properties (comparable to wrought material) as well high resolution and part accuracy. The machine deals with a range of metal powder, including Titanium-6Aluminum-4Vanadium/ELI, Cobalt Chrome and CP Grade II Titanium. The EBM A2 RCAM purchased in 2008 by SMU’s Research Center for Advanced Manufacturing is the main manufacturing mean for the aforementioned customized dental implant prototypes; in addition to the fields of automation, aerospace, biomedical, and the military (Figure 1).Figure 1: Schematic of Electron Beam Melting®, and customized dental implantThe CAD Model and the Mesh: The initial step of FEA is to produce the components of three solid CAD models. The first CAD model contains the following elements: mandibular second molar, mandibular Periodontal Ligament (PDL), mandibular cortical bone, mandibular cancellous bone, maxillary second molar, maxillary PDL, maxillary cortical bone and maxillary cancellous bone.

In the second model, the mandibular molar is replaced by a customized dental implant with a regular solid root and a porcelain crown, and the mandibular PDL is removed. In the third model, the customized dental implant has a modified outer layer of the root; the elasticity of that layer is subject to optimization in order to minimize stress shielding. All three CAD models have similar geometrical features and dimensions (Figure 2).In order to produce the dentitions and the corresponding customized dental implants, the digital acquisition of the shape of the natural tooth is needed. Each of the three CAD models goes through a similar set of data acquisition and data processing steps, those steps can be summarized as follows: (Table 1 )A CT scan of an anonymous patient, composed of 200 images interspaced by 0.1 mm is obtained. The scan covers the mandible area which displays a second molar.

The scan covers the mandible area which displays a second molar. The CT scan imported into Mimics of Table 1: Method for obtaining the CAD Model Materialise®, is processed to produce a 3D model of the second molar exported as a .stl (Standard Triangulation Language) file. The .stl file is imported into Magics of Materialise® where .stl fixing module to produce a water tight model. Next, Geomagic® Studio V.11.0 is used to produce surface manifolds patches in order to generate a solid model of the molar, which is exported as an .iges (Initial Graphics Exchange Specification) format.

And finally Autodesk® Inventor® 2010 is used to produce two implant designs, and the rest of the bio-components corresponding to the three CAD models.CT scan of a second molar in the mandibleMimics V.12 MaterialiseMagics V.12.0 MaterialiseGeomagic Studio V.11.0is employedAutodesk Inventor 2010ANSYS 11.0 Workbench 11.0andFigure 2: Schematic of the three models of the FEAA 3D representation of the first model and its corresponding mesh is shown in Figure 3. The mesh control is embedded in ANSYS® 11.0 Workbench® software, which has a useful defeaturing algorithm that enables the program to ignore sliver areas, hence producing a robust mesh and avoiding the need for CAD cleanup.Figure 3: Schematic of the CAD model, its corresponding mesh, and an excerpt of the two dentitions and the PDLMaterial Properties: All materials are assumed to be isotropic and linear elastic.

Table 2 displays a list of the materials, their corresponding mechanical properties and references.Table 2: List of Materials and the corresponding mechanical propertiesYoung’s Modulus (MPa) 20 000 350Poisson’s RatioReferenceHuman tooth Periodontal ligament (human molar) Cortical human jaw bone Cancellous human jaw bone Ti-6Al-4V Porcelain The elasticity of the modified outer-layer of the root in the third model is adjusted using an optimizing scheme called design modeling in ANSYS® 11.0 Workbench®.

As a preliminary step an initial Young’s Modulus of 10 000 MPa and a Poisson’s Ratio of 0.33 has been selected.0.30 0.45Abé et al (1996) [4] Atmaran and Mohammed (1981) [5] Abé et al (1996) [4]20 0000.330000.31Abé et al (1996) [4]110 000 689000.33 0.28Colling (1984) [6] Lewinstein (1995) [7]Boundary Conditions and Loads:A static analysis of the three models us executed, where a static load is applied, and static entities such stress, strain and deformation are evaluated. ANSYS® Workbench® has the ability to detect contact regions. The three models have respectively 7, 8 and 9 contact regions. In this simulation, the nature of contact is selected to be bonded.A human bite load is in the range of 200 N to 500 N [8].

The three models have respectively 7, 8 and 9 contact regions. In this simulation, the nature of contact is selected to be bonded.A human bite load is in the range of 200 N to 500 N [8]. A pressure corresponding to a load of 400 N is applied on the top surface of the cortical bone of the maxilla and fixed support is applied on the bottom surface of the cortical bone of the mandible (Figure 4).Figure 4: Boundary Conditions and Loads of the 1st modelDiscussion of the Results of the first two models: The time needed to solve the three models was in the order of a few minutes. The very short lead time of obtaining results using FEA is one of the main advantages of technique as well as cost-effectiveness, especially compared to lengthy and costly in vitro and in vivo experiments.Examining the results of the first model (Figure 5), a maximum deformation of 23.5 microns is observed at the bottom area of the cancellous and cortical bone.

This value of deformation is under the destructive value of 150 microns [10]. The deformation within the natural tooth is in the range of 7 to 13 microns. A few probes are positioned in the cancellous bone close the PDL, and corresponding probes are selected across the PDL within the dentition. The stress range of the probes in the cancellous bone is between 1.3 to 3.6 MPa, which matches the stress range of 1.4 to 5 MPa suggested by Rieger et al. [9] for optimal bone health.a)b)Figure 5: Plot of the total deformation (a), and the equivalent stress in the first model (b)The stress range of the probes selected within the tooth is 2 to 6.8 MPa. The stress in the tooth is slightly higher than the stress in the surrounding cancellous bone. This difference of stress ranges is an important indication of the occurrence of stress shielding.

The stress in the tooth is slightly higher than the stress in the surrounding cancellous bone. This difference of stress ranges is an important indication of the occurrence of stress shielding. The PDL in the first model is considered a linearly elastic material, which is a simplification compared to the commonly known hyper-elastic and visco-elactic portrait of the material [11].In the second model, the lower jaw molar is replaced with a customized dental Implant, and the implant/bone interface is chosen to be a direct contact interface. By applying the same loads used in the first model, the range of stress at the probes selected in the cancellous bone at the vicinity of the interface is 0.4 to 2.5 MPa, which is lower than the range obtained in the first model (Figure 5, b). In comparison, the range found in the probes selected in the customized dental implant is 2 to 9.4 MPa.

In comparison, the range found in the probes selected in the customized dental implant is 2 to 9.4 MPa . The stress difference between the implant and the surrounding bone is larger than the stress difference found between natural dentition and bone, which suggests a stress shielding effect. The maximum deformation is 17 microns (Figure 5, a), which is close to the range of optimal micromotion for Osseointegration of 30 microns [12] and below the destructive value of 100 microns suggested by the rule of thumb of Brunski [13].a)b)Figure 6: Plot of the total deformation (a), and the equivalent stress in the second model (b)Third Model and Design Optimization: In the third model, a modified outer-layer surface with adjustable elasticity is incorporated to the root of the implant. Such elasticity is be obtained and adjusted by either the use of lattice structure design or porous structure design in the production process.

Such elasticity is be obtained and adjusted by either the use of lattice structure design or porous structure design in the production process. Furthermore, the elasticity of the surface can be controlled in order to better optimize the biomechanics of the implant. The third set of results (Figure 7) is an example of the biomechanics when the elasticity of the surface is set to 10000 MPa.a)b)Figure 7: A plot of the total deformation (a), and the equivalent stress in the third model (b)It is noticeable that the stress shielding is reduced, and the micromotion is slightly increased, which implies that controlling the elasticity of the outer-surface does indeed affect the deformation, and the load transfer when compared to the second model (where the modified outer-layer is non-existent). These findings led to the use of an optimization module called DesignXplorer in ANSYS® 11.0 Workbench®, which uses a deterministic method of changing design variables to find an optimal outcome.

These findings led to the use of an optimization module called DesignXplorer in ANSYS® 11.0 Workbench®, which uses a deterministic method of changing design variables to find an optimal outcome. Central Composite Design (CCD) is the underlying Design of Experiment technique selected to obtain higher accuracy in the experimental design.Figure 8, Schematic of the locations of the 15 probes inserted into the third modelReferring to Figure 8, five probing areas are selected on a transversal cross section. Every area contains three stress probing positions; on the cancellous bone, the modified outer-layer and the implant solid core. Computing the variance of the values of stresses in each group of three probes provides a qualitative estimate of the stress shielding at that location. The greater the variance of stresses, the larger the stress shielding, and vice versa. Variance is the measure of the average deviation of the elements of a population from their mean values.

The greater the variance of stresses, the larger the stress shielding, and vice versa. Variance is the measure of the average deviation of the elements of a population from their mean values. The square variance of a population is definedby: 𝜎2 = value.𝑛 𝑘=1(𝑥𝑘 −𝜇 )2 𝑛, where n is the number of elements, xk the element with index k, and µ the meanThe input parameters are selected to be the biting pressure applied on the model and the elasticity of the interface. The parameters’ corresponding ranges are 0.25 to 2 MPa and 5 000 to 200 000 MPa respectively.Table 3, Optimal Elasticity at the five probing areasProbe AreaOptimal Elasticity (GPa)1 1002 1305 4 110 120 843Figure 8Figure 9 (1-5) displays 3D plots of the variance versus the pressure applied, and the elasticity of the layer. As the pressure applied increases the variance increases. While in the range of elasticity, a local minimum value of variance can be detected at every single probe area.

As the pressure applied increases the variance increases. While in the range of elasticity, a local minimum value of variance can be detected at every single probe area. Table 3 provides a summary of the optimal elasticity found for different probe areas. An optimal elasticity corresponds to a minimum variance, hence minimum stress shielding.This final finding is applicable to the ability of prototyping technique based on Electron Beam Melting® to produce parts with high geometric complexity and locally tailored mechanical properties. By finding the desired elasticity at different areas in the implant, the design of the porous structures is reciprocally optimized and manufactured in order to provide the desired elasticity.Figure 9 (6) displays the aggregate variance of the five probing areas.

From a global perspective, it is still very eminent that an optimal elasticity is evaluated based on the minimal variance of stresses, hence minimal stress shielding.1)2)3)4)5)6)Figure 9: Plots of the Stress variance at the 5 corresponding probing areas, versus the selected ranges of pressure and interface elasticity (1-5), Plot of the aggregate variance of the five probing areas (6)Conclusion: In conclusion, FEA provided numerous improvements by optimizing the design through the evaluation of the biomechanics of natural dentition and customized dental implants.

The model displayed in the current research serves as a foundation for more complicated models which aim towards a more sophisticated representation of the implant/bone interface and the PDL, which provides more accurate and more realistic results.Furthermore, the current exhibits the benefit of rapid manufacturing in providing prominent design flexibility, which is embodied by the concept of customized dental implants and the various geometrical and mechanical modifications it offers in comparison to traditional standardized dental implants. These modifications contribute to faster healing time, better implant retention and enhanced patient comfort and satisfaction.Acknowledgement: The US Army Research Office granted SMU a grant to develop a program entitled: “Advanced Precision Supply Parts Manufacturing”, Grant No. W911NF-07-2-0056.

sign Optimization Strategy for Multifunctional 3D PrintingA. Panesar, D. Brackett, I. Ashcroft, R. Wildman and R. Hague Faculty of Engineering, University Park, University of Nottingham, NG7 2RD, UKAbstractAn optimization based design methodology for the additive manufacture of multi- functional parts (for example, a structure with embedded electronic/electrical systems and associated conductive paths) is presented. This work introduces a coupled optimization strategy where Topology Optimization (TO) is combined with an automated placement and routing approach that enables determination of an efficient internal system configuration. This permits the effect of the incorporation of the internal system on the structural response of the part to be taken into account and therefore enables the overall optimization of the structure-system unit. An example test case is included in the paper to evaluate the optimization strategy and demonstrate the methods effectiveness.

An example test case is included in the paper to evaluate the optimization strategy and demonstrate the methods effectiveness. The capability of this method allows the exploitation of the manufacturing capability under development within the Additive Manufacturing (AM) community to produce 3D internal systems within complex structures.1 IntroductionThis paper presents and evaluates an optimization strategy for the design of multifunctional components to be made using Additive Manufacturing (AM) multi-material processes. By definition, a multifunctional component must have multiple uses, such as structural and electrical functions, e.g. a Structural Health Monitoring (SHM) component. While manufacturing processes capable of physically realizing these components are still under development, a variety of techniques have been proposed, primarily using stereolithography and direct write/print technologies. The reader is directed to [1] for a history of work carried out in this area.

The reader is directed to [1] for a history of work carried out in this area. The EPSRC Centre in Innovative Manufacturing in Additive Manufacturing at the Universities of Nottingham, UK, has the development of multi-functional 3D printing processes as one of its main aims. This Centre aims to achieve this is via multi-material jetting.While work has been ongoing within the AM community to develop the manufacturing processes to achieve Multi-Functional AM (MFAM), there appears to have been little effort to develop design philosophies/tools tailored to exploit the design freedom associated with MFAM, hence, the Centre also focuses on developing design and analysis methods to enable this. The motivation for this work lies in the realization of an ultimate aim which is to be able to intelligently optimize the design of a multifunctional part, particularly within the scope of optimal placement and routing.

The optimized 3D placement of internal components and associated routing would enable more compact, better integrated and capable MFAM systems.

This concept is illustrated by the examples in Figure 1.(a)Internal system (routed connections / wirings)Internal system (placed components)(b)(c)Figure 1: Multi-material jetted concept prototype - a) an example of a topologically optimized structural part with integrated internal system of placed components and the associated routing, b) a prosthetic arm with embedded systems and the associated routing [2], and c) a zoomed-in view of the prosthetic arm highlighting the complexity of the hand.Imparting such functionality to a MFAM part can be best achieved by coupling a placement and routing optimization with a Topology Optimization (TO) routine which is a structural optimization technique that iteratively improves the material layout within a given design space, for a given set of loads and boundary conditions [3][4][5][6].

Automated placement and/or routing techniques have been employed in numerous fields ranging from: electronics, civil, aerospace, navigation systems, and artificial intelligence (robotics). In principle it would be best to perform placement and routing in one step as placement has significant repercussions on the routing but due to the nested dependencies these can be more efficiently (in terms of computational expense) tackled independently. To this end, several graph algorithms and mathematical methods, as reported in [7][8][9], have been developed and implemented. Many of these strategies have been adapted and coupled with global optimization algorithms such as Genetic Algorithms (GA) and Ant Colony Optimization (ACO) to solve optimization problems in other fields.

Examples include pipe/cable routing [10][11][12][13] and optimum placement for SHM [14][15].To facilitate the attainment of the overarching aim, which is to optimize the design of a multi-functional part, this work presents a coupled design optimization strategy. The paper takes the following structure: firstly, the strategy for optimization of multifunctional design is outlined; secondly, the details of the placement and routing approaches, and the coupling strategies are discussed; and thirdly, the appropriateness and effectiveness of the strategy is demonstrated by evaluating and discussing the results for an example test case.2 Strategy: Design optimization for multifunctional 3D printingFigure 2 shows a coupling between a TO routine and a placement and routing optimization.

This coupled optimization strategy is essential to exploit the design freedoms offered by MFAM.Figure 2: Strategy for coupling placement and routing with topology optimization.In previous work [16], the authors demonstrated the capability of the aforementioned coupled optimization tool for a 2D test case of SHM. This preliminary work looked at integrating the placement and routing methods into a structural TO algorithm so that the optimization takesaccount of any effect that the placed components and circuitry has on the structure and makes according topological modifications. In this paper, we present the extension of this capability to 3D. Figure 3 shows an example that adopts the framework of Figure 2 wherein we have a one-to- one communication between the structural analysis and placement and routing optimization.

Figure 3 shows an example that adopts the framework of Figure 2 wherein we have a one-to- one communication between the structural analysis and placement and routing optimization. Note that the solid part of the structure is removed to allow the internal routing and components to be viewed.Figure 3: An example demonstrating the coupled optimization strategy of Figure 2.3 MethodologyIn order to exploit the increased design freedom offered by MFAM design, strategiessuitable for: 3D placement and routing, and its efficient coupling with TO need to be devised.3.1 Voxel modelling environmentA voxel modelling environment (where a voxel represents a cube in space) was considered for this study. This choice enables direct mapping to the raster based file formats used in AM, such as the bitmaps used in jetting.

This choice enables direct mapping to the raster based file formats used in AM, such as the bitmaps used in jetting. This eliminates the need for manual CAD operations, including conversion to the common STL file format and associated slicing, which is well known to be cumbersome and error prone. In addition, working in the voxel environment offers great flexibility as it enables in simple mesh mapping between different stages of the process to allow different modelling resolutions for control over accuracy/detail. This mesh mapping can be best achieved through the use of foreground meshes that are compatible with (i.e. an offspring of) a constant background mesh. We term this multi-resolution mesh philosophy as the Multiple Compatible Mesh Method (MCMM). To illustrate the usefulness of the MCMM, consider an example where a fine resolution is used for structural optimization, a coarse resolution for placement and routing optimization, and a very fine resolution for manufacturing.

In doing so, we make the overarching design, optimization and manufacturing process, more efficient.3.2 Placement and routing methodologyFigure 4 provides a top-level description of the proposed placement and routing methodology. Placement of the component involves: identifying potential locations; identifying the orientation for the component under consideration; and finally assessing the location suitability for this component. Routing involves: separating the component connection by type; computing shortest paths for pairs of components as described in Figure 4; and finally solving the combinatorial network problem (if one exists).Figure 4: Description of the placement and routing strategy.3.2.1 Skeletonization: characteristic information regarding the geometrySkeletonization is the general name given to a process which reduces the quantity of geometric information (i.e. dimensionality) required to represent a structure whilst preserving the essence of the topology.

dimensionality) required to represent a structure whilst preserving the essence of the topology. In 3D, this means a 2D medial surface and a 1D medial axis. A thinning algorithm, as detailed in [17][18], has been used to obtain the skeletal information of the part’s topology. For this study, the medial axis is of particular importance as it is used to obtain appropriate orientations of placed components.3.2.2 Placement strategyIn order to automate the component placement, we propose a placement strategy that, in principle, capitalises on both the performance and geometric aspects by coupling them. These two aspects have been outlined in Table 1 and are discussed in-depth in the following sections.Table 1: Geometry and Performance based placement strategyMedial axisGeometry Based VolumeSpecifiedPerformance BasedStructural responseOn other physics (e.g. thermal)e.g. when good component encapsulation is desired and/or a measure of member deflections is needed.e.g.

thermal)e.g. when good component encapsulation is desired and/or a measure of member deflections is needed.e.g. thermocouples may be placed at high temperature regions.1. Performance characteristics – for many structures, the internal system of components and sensors can be used to provide some assessment of the structures’ performance in-service. For example, consider Figure 5, which shows a structure subjected to an external heat stimuli. In order to effectively monitor the thermo-structural response of this structure, two thermal sensors are placed at the hot spots, a Central Processing Unit (CPU) in the cold spot (to process data from all sensors) accompanied by a thermal sensor to monitor the temperature of the CPU and four strain gauges in the key structural members.Figure 5: An example demonstrating component placement based on performance measures.2.

Geometric features – for many structures, the internal system of components and sensors might be required to be placed such that they are well encapsulated within the structure and also monitor the structural performance of the geometric members. Another example of geometrically important part location is an external sensor, such as an optical sighting or measurement device. The components could, therefore, be placed based upon the skeletal information of the geometry such as the medial axis. Doing so reduces a 3D volume to a network of 3D lines upon which the placement can be based, thereby significantly reducing computational expense. In addition to the aforementioned approach, one could adopt a more generic, unconstrained approach where the whole design volume can be exploited for placement.

In addition to the aforementioned approach, one could adopt a more generic, unconstrained approach where the whole design volume can be exploited for placement. This would be computationally more expensive as it operates on the whole 3D structural volume so it is recommended to identify a set of somewhat uniformly distributed random placement locations within the volume of the part to use for component placement. The procedure for doing this involves initially selecting a set of random points within the part volume, constructing a Voronoi decomposition from them, and then identifying the centroids of these regions which define the new point set.The approach adopted in this study for the placement of components is summarized bythe flowchart of Figure 6.

Details of this approach can be found in [19].Figure 6: The placement methodology3.2.3 Routing strategyOnce the internal components have been placed, the next task is to generate the connections to form a circuit, commonly termed routing. The routing optimization aims to improve the circuit efficiency by lowering resistance, which is proportional to the conductive track length. This is, in principle, achieved by identifying the shortest paths between components subject to design rules and constraints. By doing so, we also minimize the utilization of the conductive track material.The main constraints imposed on the routing optimization were to avoid obstacles (e.g. internal components and void regions) and to have a minimum spacing between routes (to avoid electrical interference).

internal components and void regions) and to have a minimum spacing between routes (to avoid electrical interference). Control of the track diameter was also incorporated into the method to ensure the required levels of conductivity and insulation could be achieved.A MATLAB implementation [20] of the Fast Marching (FM) method [21] was used to identify the shortest route between any two considered components and an in-house implementation of Ant Colony Optimization (ACO) [22] was used for solving the combinatorial network problem (as described in Figure 4). The robustness of ACO has been well documented [22]. Furthermore, benchmarking of the in-house ACO code was carried out using standard TSPs taken from [23][24] for which the minimum tour lengths were known and the implementation enabled in converged solutions within reasonable iterations (i.e.

computational effort).3.3 Coupled optimization procedureAs discussed by the authors previously [16], it is beneficial to first establish an initial approximation to the topology by using sufficient TO iteration until a particular criterion was met (e.g. x number of iterations or x convergence tolerance met) and only then coupling the TO with the placement and routing optimization. Herein, a one-to-one communication between the TO and placement and routing optimization is considered. It is proposed that following each structural optimization iteration, the placement of the components is determined, associated routing is performed, and subsequently the design variables are updated accordingly for the next iteration of the TO phase, however, there are numerous variations to this overall philosophy that may be preferred in certain applications.

This approach is summarized in Figure 7.Initial structural topology optimizationDefine optimization modelCircuit integration criterion met?YesComponent placementNoStructural analysisRouting between componentsFilter element sensitivitiesUpdate design variablesNoConvergence achieved?YesTerminateFigure 7: Flowchart showing the coupled optimization procedureFor the structural analysis model, the Young’s modulus of the elements representing the structure, void, placed components and wiring were updated in accordance to Table 2. Here, and are parameters that govern the property of the components and wiring by bounding them to those of the structure and void.

Here, and are parameters that govern the property of the components and wiring by bounding them to those of the structure and void. A commercial Finite Element Analysis (FEA) solver [25] was used to perform the structural analysis.Table 2: Material properties used for structural analysisFunction Structure Void Components WiringYoung’s Modulus, E 1 1E-4 ( ) ( )Poisson’s Ratio, v 0.3 0.3 0.3 0.33.4Improving computational efficiencyTo allow for an efficient optimization procedure, approximations were made without compromising the accuracy at the FEA stage. This was achieved by utilizing a hierarchical mesh decomposition strategy, specifically quadtree (2D) or octree (3D) decomposition (see Figure 8a), which consisted of three key concepts.The first of these concepts is a dual mesh system, which allows decoupling of the design variables (background mesh) and the analysis (foreground) mesh to enable flexibility of the mesh configuration.

This background mesh was of constant predefined resolution and remained unchanged throughout the optimisation. At each optimisation iteration, the analysis results from the foreground mesh were mapped onto the background mesh prior to the stage of updating the design variables. At each iteration of the optimisation a completely new mesh was generated based on the aforementioned conditions using the updated design variables.The second concept is the decomposition procedure which generated the adapted analysis elements from the design variable mesh. The primary condition used to determine when an element should be split was if there was some level of heterogeneity in the corresponding area of the background mesh. Constraining this decomposition were three primary parameters, namely, the minimum and maximum element dimensions and the maximum number of adjacent elements to any one larger element edge (2D) or face (3D).

The differing effects of these parameters on the efficiency and efficacy of the analysis and optimisation stages are discussed and recommendations made for parameter combinations are given in [26].The third concept is the use of multipoint constraint (MPC) equations to allow the use of hanging nodes in a mesh with undistorted elements by defining their relationship to their surrounding reference nodes.

This approach enables a relationship between the foreground and background meshes to be maintained, which is a necessary requirement for the Multiple Compatible Mesh Method (MCMM).a)b) Generate foreground meshSolve FEA modelRead element strain energy resultsCalculate strain energy density and map onto background meshApply checkerboard filter to background meshUpdate design variables (void-solid) using background mesh Figure 8: a) An octree decomposed topology used for analysis, b) Integration of meshing procedure into BESO TO routine.The revised bi-directional evolutionary structural optimisation (BESO) method [27][28] was used for this work, primarily, because it made identification of boundaries straightforward as the structure is inherently discrete throughout the optimisation (due to the two possible design variables values).

As a non-uniform mesh was used for this work, the sensitivity number included the effect of the element size; the strain energy was mapped onto the background mesh elements after dividing by the foreground element area or volume to get the strain energy density [4]. The rest of the procedure is typical for the BESO methodology from the aforementioned sources. The overall optimisation procedure used for this work is summarised by the flowchart in Figure 8b.4 Simulation Results and DiscussionThe test case of Figure 9 is considered to evaluate and better understand the proposed coupling strategy. In this example, the structural criteria of Figure 3 (i.e. same boundary and loading conditions) with four internal components that are all connected to one another is considered. The placement of the components is chosen to be on the mid-points of the medial axis members (refer [19] for details).

The placement of the components is chosen to be on the mid-points of the medial axis members (refer [19] for details). It was decided to incorporate the automated placement and routing into the coupled optimization at 50th iteration and terminate the coupled optimization at 70th iteration. This was done so that the structure was reasonably well-defined before considering the internal system integration.The results provide an insight into the strategy employed for coupling of the TO with the placement and routing approach. It is evident from the evolution of the solution that the proposed strategy is appropriate for the design of MFAM parts but it is the effectiveness of this strategywhich is in question. This strategy has successfully demonstrated the capability to automate the placement of components and generate the optimal routing.

This strategy has successfully demonstrated the capability to automate the placement of components and generate the optimal routing. Nevertheless, more work needs to be done to improve its effectiveness as the internal system configuration was found to be erratic despite the objective function (i.e. total strain energy) being reasonably stable. The reason for this erratic behaviour is principally because of the criterion used for placement (midpoint of medial axis members). To overcome this, two procedures can be adopted. Firstly, a system to track the placed components could be used which should allow for more stable internal system configuration changes as the optimization progresses.

Firstly, a system to track the placed components could be used which should allow for more stable internal system configuration changes as the optimization progresses. Secondly, the internal system could be introduced gradually to the structure with the values of and (in Table 2) slowly varying from 0 to 1 over the iteration history (unlike the constant value chosen for this study).a)50th iter.b)55th iter.c)60th iter.d)65th iter.e)70th iter.f)Final topologyFigure 9: Results for test case 1 – coupled optimization strategy Internally placed components and associated routing shown for a) 50th iter b) 55th iter c) 60th iter d) 65th iter e) 70th iter f) final topologyDevising a tool to facilitate the intelligent placement of components is challenging; not only is the tool required to place components at potential sites but to also place them such that the associated routing length is minimized.

Therefore the automated placement and routing methodology can be extended to include more advanced features such as: a combinatorial optimization scheme to identify the best placement configuration that minimizes total circuit length; and the capability to locally adapt the structure for a more effective routing that can result in the total route length being minimized further (this is especially important when several shortest routes pass via or would benefit passing via the same thin member).Moreover, the design freedom enabled by AM provides opportunities to create non- traditional component geometries. For example, a standard strain gauge would have predefined terminal locations. Note that in this case, these have not been predefined as these were defined by the routing of the connections to these components to avoid over-constraining the routing. More significant changes include having strain gauge coils in multiple dimensions such as those shown in Figure 10.

More significant changes include having strain gauge coils in multiple dimensions such as those shown in Figure 10. This geometry could be aligned with the longitudinal axis of structural members, particularly thin members that could not contain this number of coils in planar form. More importantly, it offers the potential for more compact and better integrated system designs.Figure 10 – Example of design freedom in creating a compact 3D strain gauge to be placed in thin structural members.5 Concluding RemarksThis paper has presented a coupled optimization strategy where a topology optimization method is combined with an automated placement and routing approach for the design and optimization of multifunctional AM parts. The proposed placement approach capitalizes on both the performance and geometric aspects of the considered part. A medial axis based component orientation scheme is proposed for an appropriate alignment.

A medial axis based component orientation scheme is proposed for an appropriate alignment. With regards to routing length minimizing problem, a FM method in conjunction with an ant colony optimization algorithm was used. Adaptive meshing, specifically, Octree decomposition is employed to obtain computational leverage in the coupled strategy. The result from the example test case demonstrated the appropriateness of the proposed coupling strategy. The capability of the placement and routing method allows the exploitation of the manufacturing capability under development within the AM community to produce 3D internal systems within complex structures. The primary next steps of this work are to devise strategies for closer integration of the placement and routing methods with the topology optimization procedure.References[1] A. J. Lopes, E. MacDonald, and R. B.

sign Rules for Additive ManufactureI. Gibson, G. Goenka, R Narasimhan, N BhatIntroductionAs Rapid Prototyping (RP) has evolved to become Additive Manufacture (AM) it has become possible to exploit the effects of the layer-based approach to fabrication so that parts can become easier to build and with greater functionality. Initially parts made in a manner that could not be manufactured further using conventional processes were considered merely as a novelty to RP machine users rather than of any wider benefit. With improvements in the technology and increasing demand for more complex geometry products, AM is coming into its own. As a result, we are now able to manufacture parts that have interlocking features, overlapping and mechanistic properties as well as us not having to worry about constraints like draft angles and overhanging features.However, this does not mean we can completely ignore manufacturing constraints.

More that we have to concern ourselves with a different set of constraints. AM technologies are limited according to the following (1): -Speed of build Accuracy Surface geometry TolerancesWall thickness and feature size Material properties Range of materialsThese have always been an issue; it’s just that now improvements have now gotten to the stage where parts are acceptable for final application. We must therefore concern ourselves with a new set of design rules that allow us to understand how the technology can be best applied to the advantage of the creator of the new products.BackgroundIt has always been known that parts could be made using (what was then known as) RP that could not be easily manufactured in a conventional manner. This was indeed what made it easier to use the technology in the early stages of the product development process. There was no need to consider draft angles or seriously worry about overhangs, radiuses, ribs, etc.

There was no need to consider draft angles or seriously worry about overhangs, radiuses, ribs, etc. that may affect the quality of a final moulded part, thus allowing the designer to focus on the form and function of the design. However, it was always considered that the final product would be made using conventional processes because the first RP machines were expensive, inaccurate and slow to build a single part from materials that were also expensive, had poor mechanical properties and suffered from long-term degradation and distortion effects.Mainly incremental improvements in the technology have focused on lowering machine costs whilst increasing build speeds and accuracies. Parallel improvements in materials have resultedin a wider range with superior properties that are consequently suited to a much wider range of applications. Many of these applications are now destined for final use rather than just earlier stages of the product development process.

Many of these applications are now destined for final use rather than just earlier stages of the product development process. Furthermore, it is now possible for example to include integral gears and cams, mechanical and living hinges, snap fasteners and even fully interlocking meshes like chain mail into a design and in a single manufacturing stage with AM technology (like the examples shown in figure 1). Metal systems have also become a reality in recent years to enhance this rich arsenal of tools that can tie design to manufacture in an unprecedented manner.Figure 1: AM feature samples (Courtesy EOS and Shapeways)It is only now that designers are waking up to the possibilities this technology can provide. Companies like Shapeways (2), Freedom of Creation (3) and MGX (4) are at the forefront of this wave, providing innovative designs that exploit the capabilities of AM technology.

Companies like Shapeways (2), Freedom of Creation (3) and MGX (4) are at the forefront of this wave, providing innovative designs that exploit the capabilities of AM technology. However, these companies are also exploiting their position in the marketplace, having extensive understanding and experience of the technologies. There are currently very few tools that support the design process with focus on AM technologies. Probably the most extensive research has been undertaken by Campbell (5), who has focused on the development of a database that indicates to the designer what different features can be included into a new product. This will enable the designer to exploit new design possibilities but falls short of quantifying the outcomes of any new design proposals.

This will enable the designer to exploit new design possibilities but falls short of quantifying the outcomes of any new design proposals. Earlier work by Campbell and by Choi (6) tried to simulate AM technology so that users and designers can understand how parts might be affected by the layer- based technology and the orientation inside the machine, for example, but this was more analytical and did not provide any initial rules for such a design. EOS carried out a study that proposed a variety of living hinge designs based on a basic set of material properties (7). This study can help designers understand how different designs might function but it is still somewhat difficult to understand how to fine-tune a design using the proposed hinge designs.When designing parts that will finally be built using AM technology, we are not talking about eliminating process constraints but incorporating a new set of constraints.

Definitely these constraints are more relaxed in terms of overall functionality of part, but we must still understand the process and materials fully in order to get the best out of the resulting parts and to avoid mistakes. The purpose of this study is to try and understand the material issues in parts made using AM so that we may be able to generate a set of quantifiable rules that can allow designers to exploit them. The ultimate aim of this work is to provide sets of rules that can allow reliable designs that also stretch the capabilities of the processes. There are numerous AM technologies available, with new processes and machines being introduced to the market at regular intervals. Existing machines are also extended by the introduction of new materials and innovative post-processing techniques.

Existing machines are also extended by the introduction of new materials and innovative post-processing techniques. It is therefore important to focus on design rules that are flexible enough to accommodate these technological developments, so that we can: -Choose the correct process and materials from what is available Define lower limits that can ensure a design feature will work from the chosenmachine parametersDetermine whether a specified design might work once it has been fabricatedOur initial study was to investigate whether we can apply a more precise set of rules to living hinge design. This study quickly moved on to a more detailed understanding of the living hinge concept so that it was possible to develop and validate numerical models that exploited in the basic principles of viscoelasticity in the materials concerned.

The rest of this paper will therefore discuss the basic mechanical properties before going on to describe a series of experiments that support a numerical model of a viscoelastic feature. We will then proceed to a discussion on how this work can assist in providing support for product designers who want to use AM technology.Study into elastomeric properties of AM materialsInitial Analysis: Elastomeric properties can be included in a wide range of applications to produce living hinges, snap fits, seals, shock-absorbers etc. While different features ideally require a unique set of material properties, most of the features which make use of the elastomeric properties require the material to able to bend, stretch or compress.

The objective of the study was to understand the kinematics of the bending mechanism which could be used in features like elastic hinges and find a suitable AM material which would be able to undergo the maximum amount of bending before failure by yield or fracture and be able to withstand a small amount of load. The best material is the one which for given ligament dimensions bends to the smallest radius without yielding or failing (8). When a ligament of thickness 2t is bent elastically to a radius R, the surface strain isAssuming that the bending is initially elastic, the maximum stress isThis must not exceed the yield or failure strength bent without damage is. Thus the radius to which the feature can beThe best material is the one that can be bent to the smallest radius, that is, the one with the greatest value of the indexIt should also be able to withstand a certain amount of axial load (F).

Assuming that w is the width and t is the thickness, the tensile stress in the feature should beSubstituting t in the earlier equation,This gives the second index which should be optimizedThe mechanical properties of some of the important AM processes were collected either through published data or from company representatives and compared using the two material indexes discussed before. The further in the upper right hand corner of the graph a dot is, the higher are the values of the 2 indexes and the better are the elastomeric properties of the material.

The further in the upper right hand corner of the graph a dot is, the higher are the values of the 2 indexes and the better are the elastomeric properties of the material. The AM materials have been compared to polyethylenes (PE) and polypropelene (PP) which are the conventional types of plastics used in features which require elastomeric properties.) a P M( E / 2 σ2.0 1.8 1.6 1.4 1.2 1.0 0.8 0.6 0.4 0.2 0.0ABSComparison of AM MaterialsPPPEFullCure720DuraForm EXPA2201DuraForm PASLA 3D SystemsFDM StratasysSLS 3D SystemsPP and PEObjetEOS0.0000.0050.0100.0150.0200.0250.0300.0350.040σ/EFigure 2: Comparison of AM MaterialsIt is noted from Figure 2 that even though PE and PP have the best elastomeric properties, some of the AM materials like Duraform EX from 3D Systems and PA2201 from EOS exhibit good elastomeric properties in terms of the 2 indexes chosen.

It should also be noted that there are other factors like feature design and molecular orientation which also play a role in the final elastomeric properties of the part have been ignored in this study. For example, some of thecritical properties of PP can be improved by enhancing orientation through optimization of the melt temperature and fill speed in injection molding (9).Experimental analysis of elastomeric properties: The Objet Eden 350 machine which uses the PolyJet Matrix 3D Printing technique was chosen for the study due to availability. Photopolymers are deposited in 0.0006 in. layers (1) by the print-head which are then cured by ultra-violet light immediately producing fully cured models without the need for post-curing. The support material, which is also a photopolymer, is removed by washing it away with pressurized water in a secondary operation.

The support material, which is also a photopolymer, is removed by washing it away with pressurized water in a secondary operation. The Eden machines can only print one material at a time but the Connex machines provide multi-material capability.FullCure 720 was chosen as the material for the investigation. It is a translucent acrylic based photopolymer material developed for Objet PolyJet-based 3D Printers. It is relatively rigid compared to DurusWhite FullCure 430 and is conventionally used to build rigid models where internal details are required (10). The material does not require any subsequent machining, has a high elongation at break and good flexural toughness (11). Palli et al (12) used FullCure 720 to successfully build integrated robotic fingers using pin joints which showed very good reliability and did not fail after one hundred thousand working cycles.Studies have shown that photopolymers exhibit changes in mechanical properties with ageing and exposure to sunlight.

This may be caused by the reaction of the radicals in the occluded sites with diffused oxygen or another reactive species (1). For the purpose of this study all the specimens were fabricated together. They were stored together for 10 days at room temperature and then tested together to offset the aging effects. There is some inherent anisotropy in the parts fabricated using the PolyJet process which may be due to layer-wise nature of the process. The specimens were fabricated in the same orientation to avoid any differences due to orientation.In order to fully understand the elastomeric properties of a material and, it is imperative to have the full stress-strain curve of the material which was not publicly available for FullCure 720. As a result, compression tests were carried out following the ASTM D695-02 Standard Test Method for Compressive Properties of Rigid Plastics.

As a result, compression tests were carried out following the ASTM D695-02 Standard Test Method for Compressive Properties of Rigid Plastics. Cylindrical specimens of diameter 12.7 mm (0.5 in) and length 25.4 mm (1 in) were prepared and tested using an INSTRON AG- 25TB uni-axial tester (Figure 3) with a 25 kN load cell at a very low strain rate of 1.3 × 10-3 s-1. The effect of friction on the top and bottom surfaces of the specimen was reduced by applying a lubricant.Figure 3: Compression testThe tests give a consistent set of results as shown in Figure 4.Compressive engineering stress vs strain curve) a P M(s s e r t S100500010203040Strain (%)Specimen 3Specimen 2Specimen 1Figure 4: Compressive engineering stress vs strain curveThe material properties were calculated from the stress-strain curves obtained from the compression tests.

After compensating for the toe region, stresses corresponding to the strain values = .0005 and = .05 were recorded (11) and the Young’s modulus was calculated using the relation:andThe Young’s modulus was calculated to be 1700 MPa. It is noted that the stress-strain curve in Figure 5 exhibits a reversal of curvature indicating a softening phase over small strains followed by a hardening phase for larger strains.Living hinges were identified as the feature to be used to study the bending phenomenon since the critical aspect in the functioning of living hinges is the material property unlike other structures gears and mechanical hinges which depend on process related properties like clearances and tolerances. The three main characteristics of the living hinge design are the recess (l) the thickness (2t) and length (L1) of the neck region.

The three main characteristics of the living hinge design are the recess (l) the thickness (2t) and length (L1) of the neck region. These dimensions are illustrated in Figure 5.Figure 5: Living hinge dimensionsSpecimens with different thicknesses of the neck region were fabricated to study their effects on the bending. A jig was developed which could be used along with a tensile micro-testing machine to bend the living hinges and record the force vs vertical displacement. It is better to use a torsion testing machine instead of a uni-axial tensile tester as one can achieve 180 degrees ofrotation through circular rotation.

It is better to use a torsion testing machine instead of a uni-axial tensile tester as one can achieve 180 degrees ofrotation through circular rotation. Unfortunately the standard torsion machine was not sensitive enough to correctly record the torque on the specimens and a micro-torsion machine was not available as a result of which the tests were carried out on the tensile micro-tester which was able to achieve a 45 degrees bending as shown in Figure 6.Figure 6: 45 degrees bending test of a living hingeThe dimensions of the test specimens were measured using the Mitutoyo digital caliper with the measurement range 0-150/0.01 mm. The average thicknesses (2t) of the specimen were 0.37 mm, 0.52 mm and 0.70 mm with standard deviation of 0.047, 0.039 and 0.028 respectively. The hinge length (L1) was maintained constant at 3 mm with standard deviation of 0.016. The top recess (l) was also maintained constant at 0.40 mm.

The hinge length (L1) was maintained constant at 3 mm with standard deviation of 0.016. The top recess (l) was also maintained constant at 0.40 mm. The height of the specimens was 2.30 mm while the breadth was 3.70 mm. A summary of the results of the tests is shown in Figure 7.Force vs Displacement for varying thickness1.5)N( e c r o F10.5Thickness of 0.37 mmThickness of 0.52 mmThickness of 0.70 mm0051015Displacement (mm)Figure 7: Force vs Displacement for varying thicknessAs expected, the force required to bend the specimen increases as the thickness of the neck region increases. There is a decrease in the force once it has attained a peak value which could be attributed to the yielding and the resulting softening of the material.

There is a decrease in the force once it has attained a peak value which could be attributed to the yielding and the resulting softening of the material. It could also be due to geometrical softening occurring due to necking in the hinge region.Finite Element Analysis of elastomeric materialsThe objective of this part of the study was to develop a general Finite Element Analysis (FEA) model taking living hinges as an example which could be used to model different features which make use of the elastomeric properties of FullCure 720 or similar materials. The studyinvestigated the high deformation which occurs during the bending of a feature and examined the ability of FEA to predict the feature behavior by obtaining simulation results from a model that undergoes high element distortion.

The analysis included the modeling of the geometry and boundary conditions replicating the experiment described earlier, the material properties obtained from the compression tests and structured mesh elements to calculate the contact forces which were compared to the results obtained from the experiments. Abaqus/Standard was chosen for the study after examining the large deformation capabilities of various FEA software packages.The tests involved high strains and consequent material yielding which meant that an elastic- plastic model would be needed instead of a simple elasticity model. Since the compressive yield stress for FullCure 720 is 1.4 times higher than the tensile yield stress (10) indicating pressure- sensitive yielding (13), it was not possible to simulate the test using von Mises plasticity (14).

While heterogeneous deformation is closely approximated by Mohr-Coulomb or modified Tresca criterion, the Drucker-Pragar yield criterion is suitable for polymers which undergo homogenous deformation (15). Even though the PolyJet process exhibits a small amount of anisotropy due to the layer-wise fabrication process, the anisotropy is small as compared to other processes and not well documented. Thus FullCure 720 was assumed to be strain-rate independent and undergo homogenous deformation and the original linear Drucker Pragar model with a symmetric stiffness matrix was used with σc/σt as 1.4 in conjunction with the linear elastic model.The angle of friction is given by: (16)where α = σc/σt.was calculated to be 26.5°.The boundary conditions as shown in Figure 8 and 9 were defined to closely resemble the experiment conditions.

A part right-hand side was constrained so it could not move in the x and y direction neither rotate about the z axis (referred to as ZASYMM constraint in Abaqus) to replicate the clamping mechanism used in the experiment where a screw pressed on the feature to prevent it from moving. A rigid analytical body was used to model the load pin which was made out of steel as the stiffness of steel is much higher than that of FullCure 720. The rigid body was constrained in the directions except translation along the y axis. As the load pin moved up along the y axis it pushed the feature thus bending it.Figure 8: Boundary conditions during the experimentFigure 9: Boundary conditions in the FE modelThe Poisson’s ratio is taken to be 0.3 as published by Vesenjak et al (17).

Since Abaqus accepts only true stress-strain curves, the true stress and true strain were calculated from the engineering stress-strain curve using the relations:The subscripts t and e refer to true and engineering respectively. The model was unable to provide converging results for the softening phase in the stress-strain curve. Since the objective of the study was to model the feature behaviour at large deformations and the softening takes place over a relatively small region, the piecewise power law hardening law was used to approximate the stress-strain curve in order to get convergent solutions. Using Eqns:where assumed to be 50.

Using Eqns:where assumed to be 50. The resulting approximation is shown in Figure 10.,is the yield strain,is the yield stress and n is a constant which isApproximation of Stress vs Strain curve) a P M(s s e r t S100 80 60 40 20 00510152025Strain (%)Engineering StressTrue StressApproximationFigure 10: Approximation of Stress vs Strain curveSince the feature modeled has a constant geometry through-out the cross-section, a two- dimensional model was initially used in order to reduce the solver time. Simulations were carried for both plane stress (referred to as CPS4 in Abaqus) and plane strain (referred to as CPE4H in Abaqus) elements but the results were not accurate. While the simulations with the plane stress element type gave the correctly predicted the peak force value, there was excessive local distortion in the neck region which led to unloading in the rest of the feature.

On the other hand the simulations with the plane strain element type significantly overestimated the peak force value. Consequently a three-dimensional model was developed with 5796 eight-noded linear brick, hybrid elements (referred to as C3D8H in Abaqus). Analysis was initially carried out using reduced integration but it led to hour-glass formation and consequently hybrid formulation was adopted. The mesh was well refined near the region where the bending takes place with 75 nodes over a 3 mm region. The contact between the load-pin and the specimen was modeled using a penalty friction formulation with a frictional coefficient of 0.2 (18).The results obtained from the experiments were compared against the finite element analyses using the output variable CFN signifying the total force due to contact pressure in Abaqus.

Figure 11 shows the comparison between the specimens of average thickness 0.37 mm and the corresponding FE results while Figure 12 shows the comparison between the specimens of average thickness 0.70 mm and the corresponding FE results. It is noted that while the results agree closely for low strains, there is a slight divergence between the curves for large strains.

It is noted that while the results agree closely for low strains, there is a slight divergence between the curves for large strains. This may have been due to the piece-wise power law approximation for large strains and also the predomination of non-linear visco-elasticity at higher strains leading to a decreased Young’s modulus (19).Specimen of thickness 0.37 mm0.5)0.4N( e c r o F0.30.2Experimental Result0.10Finite Element Analysis051015Displacement (mm)Figure 11: Comparison between experiment and FE results for thickness of 0.37 mmSpecimen of thickness 0.70 mm1.5)N( e c r o F10.5Experimental Result0Finite Element Analysis051015Displacement (mm)Figure 12: Comparison between experiment and FE results for thickness of 0.70 mmFigure 13: Screenshot of the FE model on AbaqusDiscussion and ConclusionsWe appear to have a good FE model that approximates and correlates well with the experimentation.

The unloading phase which involves the visco-elastic phase could be investigated and modeled in the future. Moreover, the modeling of material softening as well as crack propagation and fatigue could lead to better prediction of failure.Based on the initial analysis with the material indexes discussed earlier, FullCure 720 looked to be a good candidate for the fabrication of living hinges. However more precise models indicated problems. FullCure 720 does exhibit elastomeric properties but not strong enough to withstand heavy use. Further studies will investigate other materials and applications.Bibliography1. I. Gibson, D. W. Rosen, B. Stucker. Additive Manufacturing Technologies, Springer , 2010.2. Shapeways . [Online] [Cited: July 7, 2010.] www.shapeways.com.3. Freedom of Creation . [Online] [Cited: July 7, 2010.] www.freedomofcreation.com.4. MGX. [Online] [Cited: July 7, 2010.] www.materialise.com/MGX.5. Implications on Deisgn of rapid manufacturing.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesign rules for additively manufactured wrist splintscreated using design of experiment methodsS.KELLY1, A.M.J. PATERSON1 and R.J.BIBB1Loughborough Design School, Loughborough University, Loughborough, UKS.Kelly@lboro.ac.uk, A.M.Paterson@lboro.ac.uk, R.J.Bibb@lboro.ac.ukAbstract: Research has shown that wrist splints can be made using Additive Manufacturing (AM) with a similar or greater performance than splints created using traditional manufacturing methods. By using AM, many of the problems associated with traditional splinting such as poor aesthetics and poor ventilation could be mitigated. However, work to date typically reviews splints with singular pattern designs (e.g. Voronoi patterns), which have structural and safety implications if similar but untested patterns are created.

However, work to date typically reviews splints with singular pattern designs (e.g. Voronoi patterns), which have structural and safety implications if similar but untested patterns are created. Using Design of Experiments (DOE) design rules were to enable clinicians to confidently design splints alongside their patients.Design rules were created by investigating variables of cut out patterns using DOE methods. Finite Element Analysis (FEA) of various combinations of cut out variables was conducted.Keywords: Additive Manufacturing, Wrist Splints, Finite Element Analysis, Design ofExperiments1. IntroductionWrist splints can be recommended for a number of wrist pathologies including wrist trauma, such as fractures or sprains, or long term conditions such as arthritis or Carpel Tunnel Syndrome (1). The aim of the splint in these cases is to provide protection and stabilisation to the affected joint, and in some cases correction or prevention of deformities (2).

The aim of the splint in these cases is to provide protection and stabilisation to the affected joint, and in some cases correction or prevention of deformities (2). Wrist post- surgery recovery can also benefit from splints to assist with rehabilitation by promoting movement in a controlled manner or through restricting movement of other more mobile joints (1).Most custom-made wrist immobilisation splints (Figure 1) are fabricated using LowTemperature Thermoplastic (LTT).853Figure 1: Traditional, custom-made wrist immobilisation splints (3)This traditional splinting process can be very time consuming, and does not always guarantee a satisfactory splint.

Furthermore, patient adherence is often a problem (4,5), sparking the need to improve the process and subsequently the outcome.There are many ways that Additive Manufacturing (AM) can be used to improve the quality of life for someone who needs to wear a splint including making splints easier to clean, more aesthically pleasing and more breathable by utilising lattice type structures, and using scan data could allow for a perfect fit.Whilst there have been a number of examples of AM splints (6–9) which have proposed improved aesthetics and ventilation in recent years, there are currently no known design rules specific to splinting which can help guide clinicians to design safe, reliable splints with minimal displacement under reasonable load for AM. It has been proposed by Paterson (10) that the traditional splinting process be replaced with a digitised method.

It has been proposed by Paterson (10) that the traditional splinting process be replaced with a digitised method. By doing this, the advantages possible with AM as discussed above could be utilised and the manufacturing process taken away from the clinician, allowing them to spend their time and their expertise to design the splint with the patient creating a sense of co-design; this is a strategy (i.e.

co- design) developed to increase patient adherence (11).Figure 2: Sketches showing the possiblities to customise the lattice structure [Image courtesy of Richard Bibb2009]Giving the power of design to patients should lead to greater compliance but leaves the clinician unable to account for all of the new variables within the design process and therefore unable to guarantee a structurally sound splint.Although previous work has shown that AM is a real solution to some of the problems discussed previously, most of the work (12–14) consist of testing on one singular cut out pattern. The process of digitising the splinting process is one that requires many aspects to change drastically, from data capture to creating a computer program to allow the clinician to854design the splint, to selecting the right AM process.

This research focuses solely on the relationships between the cut-out pattern of a splint and the corresponding displacement under loading.To continue to bring digitised splinting closer to realisation two studies investigating various factors and their effect on the displacement response have been conducted. The first study screening for significant factors and the second optimising factors, both studies have been conducted using FE (finite element) analyse and DOE (design of experiment) methodology.As it is expected that variations in cut out patterns of an AM splint have a direct impact on the mechanical properties of a splint, the DOE method was used to identify optimal cut out pattern parameters. The DOE methodology has been applied widely for optimising parameters in many fields.2.

The DOE methodology has been applied widely for optimising parameters in many fields.2. AimThe aim of this paper is to generate design rules that allow for the creation of splints, with minimal displacement when loaded, through a digitised splinting process for production using Additive Manufacturing. This will be achieved by identifying factors that are significant to the displacement response. A screening study will identify ideal factors and these factors will be optimised in an optimisation study.3. Material and MethodsOnly static wrist immobilisation splints will be considered for this study as they are the most commonly prescribed splints (15) for a wide range of conditions including a sprained wrist, rheumatoid arthritis and carpal tunnel syndrome.

This type of splint was also chosen as it is the most simplistic splint type biomechanically.3.1 FE ModelAs custom-made splints are made to fit each individual patient the topography (overall form) of the splint can vary significantly; this in turn would result in varying results with little/no chance of generalisability across testing. Therefore, it was decided to use a simple overall geometry of a cylinder to represent a splint. This helps to remove any variability from the analysis that would come from the bespoke shape of individual splints.This cylinder was created using anthropometric data as a guideline (16). Traditional splinting dictates that a splint is two thirds the length of the forearm (17). Anthropometric data was used to gather the length of an average female palm and forearm. This was used to create the dimensions of the cylinder used; 200mm length, 25mm inner radius, and wall (i.e.

Anthropometric data was used to gather the length of an average female palm and forearm. This was used to create the dimensions of the cylinder used; 200mm length, 25mm inner radius, and wall (i.e. splint) thickness was dependant on the nature of the study conducted.The geometries used in this study were created in Rhino (Version 5, Robert McNeel & Associates, Barcelona) using the Grasshopper plugin (Version August-27, 2017, Robert McNeel & Associates, Barcelona). This was used to easily create various geometry in a semi-automatic fashion.This geometry was then imported into Solidworks Simulation for FE analysis (Solidworks 2016, Dassault Systèmes SolidWorks Corporation, Tennessee). An automatic mesh generator was used with a h-adaptive meshing method. A minimum convergence rate of five percent855was set. The element type was tetrahedral 3D solid elements.

An automatic mesh generator was used with a h-adaptive meshing method. A minimum convergence rate of five percent855was set. The element type was tetrahedral 3D solid elements. Mesh coarsening was used to reduce computational time.3.2 Material propertiesPolyamide 12 (PA12) was the chosen material and a linear elastic isotropic material model was used. This was used as it was predicted that all deformation would occur within the elastic range. Material properties were taken from a Materialise data sheet of Laser Sintered PA12. This material and process were chosen as it is one of the most likely AM process to be used in AM splinting.3.3 Boundary conditions and loading convictionsIn an effort to replicate the wrist joint, the cylinder was constrained in the “forearm” region (Figure 3). A torque was applied to the “palm” region to replicate wrist flexion (Figure 3).

A torque was applied to the “palm” region to replicate wrist flexion (Figure 3). This is the most dominant wrist movement.This torque was 3Nm, a value chosen as it is similar to half the maximum value of an average adult male (18) and is also the value used in many wrist brace testing standards (19,20). This value was deemed sufficient as a wrist splint is used as a movement deterrent, unlike sport braces where the wearer may try to get the maximum range of motion out of the splinted joint.Figure 3: FEA boundary conditions on cross section of cylinder. (Torque depicted in purple, constraint depictedin green)4. Screening StudyThe objective of this study was to investigate how variables within a cut-out pattern affect the stiffness of a splint.

(Torque depicted in purple, constraint depictedin green)4. Screening StudyThe objective of this study was to investigate how variables within a cut-out pattern affect the stiffness of a splint. A general factorial study was designed with four factors including the ratio of cut out volume to volume of material in the splint, the frequency of the cut outs, the shape of the cut out, and the configuration of the cut outs.‘Cut out shape’ refers to the shape of the perforations. This factor was chosen as this isone of the main variables patients are thought to want to change.‘Ratio’ refers to the ratio of the cylinder volume to material removed. For example, a full cylinder with no cut outs would have a ratio of 0 and a cylinder with half the material removed would be 0.5.

For example, a full cylinder with no cut outs would have a ratio of 0 and a cylinder with half the material removed would be 0.5. Ratio was considered as one of the factors to be considered as having a splint with controllable volume could help with heat distribution and make the splint more lightweight.856‘Frequency’ is the second continuous factor and refers to the frequency or number of cut outs. This factor is included in this study as a single ratio value can be created with many small cut outs or fewer large ones. Again this factor was deemed worthy of investigation as it could be helpful when trying to control heat distribution.The last factor investigated in this study is ‘Cut out Configuration’. This refers to the arrangement of the cut outs on the cylinder; i.e. what pattern they are arranged relative to the surface of the cylinder. This is to investigate if a configuration can be manipulated to control the displacement of the cylinder under loading.

what pattern they are arranged relative to the surface of the cylinder. This is to investigate if a configuration can be manipulated to control the displacement of the cylinder under loading. As per factional design methodology, a high and a low level were chosen for each continuous factor. All factors and their levels can be seen in Table 1FactLevelsorsA B C DRatio Frequency Shape Configuration2 2 3 20.16, 0.5 12, 108 Circle, Square, Diamond Linear, DiagonalTable 1: Screening Study Factors and LevelsAs per DOE methodology a high and a low value for Ratio and Frequency level was tested, these values seen in Table 1 were selected arbitrarily as a starting high and low points for this work.The three shapes levels were chosen to evaluate the greatest representation of stress concentrations in the east amount of interactions.

The circle is a shape that is widely known to have less stress concentrations, where as a square or diamond has a 90 degree corner, which is likely to cause higher stress concentrations and also both allow for more uniform struts between the cut outs.Configuration levels were linear and diagonal. These variations are depicted in Figure 4.857LinearDiagonalFigure 4: Cut out Configuration levels (purple depicts cut outs through the thickness of the splint)An experimental design was created in Minitab (Minitab® 18.1, Minitab, Inc., State College, PA) with the four discussed factors all with two levels except shape. This factor was investigated at three levels. There were no replications or blocks. The design run order was randomised.4.1 Results of screening studyThe screening study was conducted and an analysis of the factorial study was conducted in Minitab.

The design run order was randomised.4.1 Results of screening studyThe screening study was conducted and an analysis of the factorial study was conducted in Minitab. A Pareto chart was created and two factors extend past the reference line, this indicates that two of the factors chosen to examine in the study are significant; Ratio and Frequency. Using automatic backward elimination, a second Pareto chart (Figure 5) was created, further confirming that Ratio and Frequency, and their interaction are the most significant factors.858Pareto Chart of the Standardized Effects (response is Displacement, α = 0.1)Term1.725AFactor Name A Ratio Frequency B Shape C Configuration DBAB0123 4 Standardized Effect56Figure 5: Backward elimination Pareto chart of screening studyThe interaction plot showed that low ratio displays little change between other factors. Higher ratio does however have an interaction with other factors (Figure 6A-C).

Higher ratio does however have an interaction with other factors (Figure 6A-C). Shape factor levels Circle and Square have no interaction with frequency (Figure 6D), but Diamond does have an interaction with the frequency of the cut outs. Square and Diamond shapes have a high interaction with configuration (Figure 6F).Figure 6: Interaction Plots (A-F) for screening study859The main effects plot shows that Ratio has a strong positive effect on the response as expected as it is a significant factor. This can be seen in Figure 7A. Shape and configuration have very little effect on the response, as shown in Figure 7C + 7DFigure 7: Main Effects Plots (A-D)Frequency has a negative effect on the response (Figure 7B). This is an unexpected result and lead to further investigation.

This is an unexpected result and lead to further investigation. As main effects plots use the mean response the spread of the response, grouped into ratio was examined.1.6001.400Plot of Displacement verus Frequency (grouped by Ratio)1.200t n e m e c a p s i D1.0000.800l0.6000.160.2450.330.4000.4150.2000.50.0000204060 Frequency80100120Figure 8: Plot of Displacement versus Frequency (grouped by Ratio)From this investigation seen in Figure 8, it is evident that frequency does not have as large a negative effect on the response as first supposed. All ratios except 0.5 show a neutral or860slightly negative effect on the response, meaning as the frequency increases the displacement decreases. The results from the 0.5 ratio are the outliers compared to the rest of the data.

The results from the 0.5 ratio are the outliers compared to the rest of the data. These inconsistencies lie in factor combinations with large cut outs quite close together.This can be seen in the points highlighted with a red arrow in Figure 9B.ABFigure 9: A - Diamond cut out, linear configuration, 20 cut out frequency and ratio of 0.5B – Close up of small strutsThese geometries were created in Rhino Grasshopper by cutting a 2D shape on a flat plane perpendicular to the 3D cylinder surface. This creates cut outs in the cylinder that cut linearly inwards (Figure 10 A+B) not radially inwards (Figure 10 C+D) meaning that the inner strut thickness Figure 10 B is smaller than Figure 10 A.Figure 10: Cross section of cylinder with cut out, linear cut out left, radial cut out right861This can be seen in Figure 9 and these small volumes between the cut outs lead to much larger displacements, skewing the mean responses.

These inconsistencies should be addressed in future work.4.2 Key findings of screening studyOf the four factors investigated in this study, two have been deemed insignificant to the displacement response. Ratio is by far the most important factor. This aligns with the initial assumption that the more material taken away from a structure the less strength the overall geometry maintains.Shape ties in heavily with ratio and should be investigated further. Shape was deemed insignificant in the screening study, but as this study has non-uniform strut thickness and some of the trials showed much higher displacements when these inconsistent struts were too small it was deemed that shape should be investigated again with consistent struts.While the configuration was also deemed an insignificant factor, like the shape factor, it should not be discarded.

While most configurations showed no interaction with other factors and showed little pull on the response, when the configuration lead to small struts between the cut outs, the response increased severely.Frequency was deemed a significant factor in this study. As discussed previously, certain combinations of frequency and other factors caused irregularities and upon inspection frequency was not as significant as first thought. The screening study showed how any weakness around the wrist joint affects the displacement response5. Optimization StudyBased on findings from the optimisation study the objective of this study was obtain significant factors and access how these factors affect the response. Using this understanding design rules could be created.Four factors were chosen for this screening experiment; Shape, Thickness, Position, Strut. Two factors are categorical; shape and position. The other two factors, thickness and strut are continuous.

Two factors are categorical; shape and position. The other two factors, thickness and strut are continuous. As discussed in 4.2 Key findings of screening study, shape was investigated further. Shape may also be a key factor to allow patient co-design so it is essential to fully understand this factorThe screening study uses a splint thickness congruent with traditional splinting guidelines. In this study splint thickness was investigated to see if this guideline transfers to AM splints or if new guidelines are needed.Position refers to the position of the largest density of splint material. This factor could allow for a denser splint in areas such as the wrist joint where the splint needs to be stronger and also allow for a less dense, more breathable splint in other areas.

This factor could allow for a denser splint in areas such as the wrist joint where the splint needs to be stronger and also allow for a less dense, more breathable splint in other areas. This could also help reduce the weight and the cost of manufacture.862Figure 11: Example of two levels of position factorAs screening study has shown that when struts become too thin splints exhibit much greater displacement, in this optimisation study, the strut thickness was controlled and struts were consistent.As per factional design methodology, a high and a low level were chosen for eachcontinuous factor. All factors and their levels can be seen in Table 2.FactLevelsorsA B C DShape Thickness Position Strut4 2 2 2Circle, Hexagon, Square, Triangle 2mm, 4mm Forearm, Wrist 5mm, 10mmTable 2: Optimisation Study Factors and LevelsAn experimental design was created in Minitab with the four discussed factors all with two levels except shape.

This factor was investigated at four levels. There were no replications or blocks. The design run order was randomised.5.1 ResultsThe optimisation study was conducted as per methods discussed previously in section 3. Material and Methods and an analysis of the factorial study was conducted in Minitab. A Pareto chart was created and the positon factor extends past the reference line, indicating that this factor is significant. Using automatic backward elimination, a second Pareto chart (Figure 12) was created. This shows that all factors are significant.863Pareto Chart of the Standardized Effects (response is Displacement, α = 0.1)Term1.708CFactor Name A Shape Thickness B Position C Strut DBDA0.00.51.01.52.02.53.03.5Standardized EffectFigure 12: Backward elimination Pareto chart of optimisation studyA main effects plot (Figure 13) was created.

it indicated that as expected a thicker splintleads to less displacement, a splint with less material around the wrist joint causes more displacement and a smaller strut thickness causes less displacement.Figure 13: Main Effects plot of optimisation studyShape is correlated with ratio as to keep the minimum strut thickness consistent the ratio will be different for each shape. This can be seen in Figure 14; here two patterns with the same factors but shape have been depicted in 2D form.864Figure 14: Two dimensional depiction of similar patterns.

This can be seen in Figure 14; here two patterns with the same factors but shape have been depicted in 2D form.864Figure 14: Two dimensional depiction of similar patterns. Label indicates shape, sturt thickness and ratio.When the ratio is plotted on the same graph it is clear that the relationship between ratioand displacement seen in the screening study hold true.Figure 15: Main Effects Plot of Shape factor (ratio in red)An interaction plot was created and there are no notable interactions between factors inthis study.8655.2 Optimisation Surface ResponseAs this optimisation is set up in a way that position is not continuous it was decided not to investigate this factor any further in this study. A more in-depth investigation will be conducted as part of future work. As any attempt to change the shape factor will lead to greater changes in other more significant factors it was decided shape would not be investigated any further in this study.

As any attempt to change the shape factor will lead to greater changes in other more significant factors it was decided shape would not be investigated any further in this study. Strut and Thickness were investigated further in a surface response study to further understand their effect on the response.A centre composite design was conducted using methods discussed previously in 3. Material and Methods. A contour plot of the results was created and can be seen in Figure 17. From the plot it is clear a more optimal design space is possible as the optimal region of lowest displacement could be more central. The design space was shifted and another analysis was performed.

From the plot it is clear a more optimal design space is possible as the optimal region of lowest displacement could be more central. The design space was shifted and another analysis was performed. The date point circle in red in both Figure 17 and Figure 18 represents the same date point and shows how the design space was moved.Figure 16: Contour Plot of optimisation surface response866Figure 17: Contour Plot of optimisation surface responseThe contour plot Figure 18 shows a clear optimal combination of strut and splint thickness. The diagonals of the contour plot also correspond with cross sectional area (Figure 19).

The diagonals of the contour plot also correspond with cross sectional area (Figure 19). A cross sectional area of 37.5mm and over correlates with displacements less than 0.5mmFigure 18: Contour plot of optimisation surface response with cross sectional area in red (mm2)8675.3 Key finding of Optimisation StudyFrom the optimisation study it can be concluded that larger cut outs should be positionedaway from the wrist joint, with a greater density of splint material around the wrist joint.The shape factor is significant as it is connected to the ratio of the splint. Strut and thickness show a clear optimal region for minimising displacement. The ideal thickness is greater than 3mm and the ideal strut is greater than 7.5mm when the cross sectional area is greater than 37.5mm2.6.

The ideal thickness is greater than 3mm and the ideal strut is greater than 7.5mm when the cross sectional area is greater than 37.5mm2.6. Conclusion6.1 Screening StudyCut outs should be constructed in such a way that the strut thickness is uniform across the thickness of the splint.The configuration and shape of the cut outs are insignificant if strut thickness is consistent6.2 Optimisation StudyLarger cut outs should be positioned away from the wrist joint • • • Ideal thickness is 4mm ± 1mm Ideal strut is 10mm ± 3mm Ideal cross sectional area of struts is 37.5mm2 or greater7. Discussion and Future WorkAlthough AM splinting is not common place practice yet, many examples can be found including one commercial venture. As with all new technologies it is important to ensure that any medical device is as safe or safer than its predecessor. AM splints are still very untested and clinicians should be cautious when designing them.

AM splints are still very untested and clinicians should be cautious when designing them. There are still many factors to be considered, including material and skin contact, optimal strut thickness for minimal pressure on the skin, optimal splint donning and doffing methods, thermal considerations, etc.It is hoped this research and the research that will follow will allow clinicians to design splint to be manufactured using AM with a greater confidence that the splint will perform as intended and allow the clinician to create splint that gives their patients a better quality of life.Future work will include continued research into the density of splint material relative to the positon on the splint. This will include separate investigations into each of the three main sections of the splint; palm region, wrist region, and forearm region.

This will include separate investigations into each of the three main sections of the splint; palm region, wrist region, and forearm region. It is hoped that this will allow for separate design rules for each region meaning a splint could be created that is lighter and could be adjusted for optimal thermal comfort. It should also reduce the cost of manufacture as each region of the splint will be the optimal amount of material. Physical testing will be conducted to validate the methodology of this work.Future work will also need to include validating design rules for different AM process, materials as this study was only conducted using laser sintering and PA12 as the material. As current technology stands this seems like the most appropriate process and material but as AM technology changes this will need to be updated.8688. Limitations of studyThe study was only on conducted on cylindrical forms, and not applied to scan data which will generate different results.

Limitations of studyThe study was only on conducted on cylindrical forms, and not applied to scan data which will generate different results. This methodology will be tested on splint geometry from scan data.Flexion is not a true representative of the biomechanical movements of the wrist due to but due to the necessary complexity of FEA that would be needed and difficulty in identifying specific fulcrums of the wrist joint a simple cantilever type boundary condition was used. However this method is a first step towards defining mechanical properties and refinement of associated design rules which define perforations of a splint.9. AcknowledgementsFunded by Materialise, supported by Loughborough University, under the EPSRC Centrefor Doctoral Training in Additive Manufacture.References1.Coppard BM, Lohman H, Shultz-Johnson K. Introduction to Splinting: A Critical- reasoning & Problem-solving Approach [Internet]. Second Edi. Mosby; 2001.

SIGN RULES WITH MODULARITY FOR ADDITIVE MANUFACTURINGH. Jee*†, Y. Lu†, and P. Witherell†Mechanical System and Design Engineering, Hong Ik University (HIU), Seoul 121-791, South Korea†Systems Integration Division, National Institute of Standards and Technology (NIST),Gaithersburg, MD 20899, U.S.A.ABSTRACTDesign rules for additive manufacturing (AM) can help ensure manufacturability, which can be viewed as compatibility between designs and the fabrication processes that produce those designs. Additionally, such rules frequently provide direct guidelines or constraints for designing AM- destined parts. Here, we present design rules as sets of modular components and associated formalisms. Independent of context, these representations can be more easily interpreted and efficiently implemented. Given context, components are specialized to represent process-specific parameters for different AM builds and processes.

Given context, components are specialized to represent process-specific parameters for different AM builds and processes. This method of specialization enables designers to reconfigure design rules, rather than create new rules from scratch, thus preserving fundamental AM principles while supporting customization and explicit representation.1 INTRODUCTION A rule can be defined as “a prescribed guide, a valid generalization, or a standard of judgment” [1] for conduct or action. Design rules, therefore, often provide direct guidelines or constraints in relation to part designs, process parameters, and material properties. Design rules also provide a means to control costs, time, and outcomes for both the design and manufacturing phases of the product realization process [2]. This control allows a designer to verify correctness in designs prior to manufacturing.

This control allows a designer to verify correctness in designs prior to manufacturing. For instance, in semiconductor manufacturing [3], ‘design-rule checks’ (DRC) help manage various complexities associated with 1) a circuit layout, including geometric representations and 2) data during the manufacture of a working design [4]. Over time, design rules often become a basic requisite for many innovation-driven fields. One such field is additive manufacturing (AM).AM technologies provide tremendous flexibility for designers because of the wide range of complex geometries that they can produce. For this reason, design rules in AM are desired because they can provide much needed insight into manufacturability for a particular material, design, or process. That insight leads to understanding direct guidelines or constraints during process planning.

That insight leads to understanding direct guidelines or constraints during process planning. Design rules can be critical to satisfying manufacturability because they enable designers to determine the best process and material combinations for a design.The aforementioned criticality is based on how mechanical properties are determined in a final part. In AM, mechanical properties of a manufactured part are as dependent on the process specifics as they are on the raw material. To compound challenges, process/material relationships are not consistent and will differ based on both the material and process type. ASTM F2792 hasdefined seven different categories of AM processes that support layer-by-layer fabrication [5]. Variations between these processes and materials significantly compound design spaces – going way beyond complications created by basic geometries. Design rules provide a means to manage complications created by those variations.

Design rules provide a means to manage complications created by those variations. Design rules are capable of limiting design features for different process/material/parameter combinations. A given set of design rules can effectively constrain a design space. However, members of this “given set” are subject to the specifics of AM materials and processes.Variations in AM builds are created from multiple different sources; these sources include materials, processes, equipment manufacturers, and machines. In AM, design rules are abstracted from observations and correlations made between material, process, and design. Design rules derived from one data source will likely differ from design rules that are derived from other data sources. In addition to inconsistencies created by data sources, ambiguities can arise from the way the design rules are communicated and implemented. Different language and syntax can influence how design rules are constructed and interpreted.

Different language and syntax can influence how design rules are constructed and interpreted. For reasons such as these, AM design rules can become so generalized that their value to the designer becomes diluted. In this paper, we propose a formal approach to design-rule representation to manage potential ambiguities.Here, we propose the adoption of principles and formalisms that allow us to modify, extend, reconfigure, or customize generalized rules as needed - instinctively and deliberately. Formalisms provide both structure for the generalizations and a means to tailor that structure for a specific process, machine, or build.

Formalisms provide both structure for the generalizations and a means to tailor that structure for a specific process, machine, or build. Overall, this approach will 1) promote the consistent application of design rule principles, 2) mitigate ambiguities and inconsistencies that may be introduced both within and between processes, 3) provide guidelines for the generation of new design rules, and 4) establish design rules as sets of modular components and associated formalisms.In the following sections, we first review sources of variations in design rules, where dilution and ambiguities can diminish their effectiveness. We then introduce our approach for incorporating formalisms into design rules.

We then introduce our approach for incorporating formalisms into design rules. In the final stage of the paper, we will demonstrate as a case study how process-independent representations provide a basis for specific interpretations of design rules for an AM process.2 BACKGROUND AND APPROACH Process-dependent guidelines and design rules have been widely developed for AM as “further process specific restrictions, explored and summarized in a simple and intuitive way, will offer a comprehensive overview of its limitations and possibilities [6].” Many researchers have focused on developing, and thus providing, prescriptive guidelines or explicit constraints for AM-destined designs, mostly for either the process categories of ‘material extrusion (ME)’ [7,8,9] or ‘powder bed fusion (PBF)’ [10,11,12,13,14,15] as defined in ASTM F2792 [5].The development of generalized design rules to accommodate process variations has also been a focus of current research.

For instance, Adam and Zimmera proposed a generalized approach for handling geometry issues in developing compatible design rules over different AM processes. implemented comprehensive design rules by defining process-independent, First, geometrical, and standard elements. Next, they developed parametric templates compatible with most AM processes. Then, they used elements and templates to develop process-specific designtheyrules based on independent functions [16]. Gibson et al., on the other hand, developed and validated a finite element model of a viscoelastic feature in materials. This model can be used to 1) provide sets of rules defining lower limits of designs and 2) create reliable designs for a given specific AM process [17]. Many other approaches can be found in [18,19,20].From a knowledge-based view, design rules constitute a body of knowledge (BoK) that contains information about allowable design actions.

Many other approaches can be found in [18,19,20].From a knowledge-based view, design rules constitute a body of knowledge (BoK) that contains information about allowable design actions. This BoK could become “a common source of a vast amount of information collected, compiled, and established by collaboration between industry and academia [13].” Therein lays the challenge! How do we provide a common set of instructions to the users of manufacturing processes that are inherently different?2.1 An “Overhang” Example Several types of AM processes are unable to build material over spatial regions without proper support. For these processes, the unsupported material, or ‘overhang,’ must be carefully managed to preserve the integrity of the design. Management of overhangs is a common example in a set of generalized design rules. Before fabrication using particular processes, support structures are often necessary to prevent the overhang from failing.

Before fabrication using particular processes, support structures are often necessary to prevent the overhang from failing. The design rule should leave no doubt when supports are necessary to build overhangs (unless overhangs can be minimized due to changes in the build orientation [21, 22]).The rule stated above has variations. For example, according to investigations by Thomas [10], rules provided for overhangs can be in the form of ‘self-supporting radii or holes’, as in the case of self-supporting angles, or vary depending on the type of overhang, material, and process.

Exact examples of such guidelines are listed below.“The lowest angle of a flat downward-facing overhang turned out to be 45 degrees [10]” (PBF process).“The smallest radius of self-supporting curve required a 28 degrees tangent on the radii and the largest one a 40 degrees tangent [10]” (PBF process).“The smallest holes of the self-supporting type appeared to be the least accurate at 0.3mm and holes at 5mm radius and above were all within a tolerance of 0.1mm [10]” (PBF process).“The common overhang distance for all layer thickness is approximately 0.075mm to 0.08mm [10]” (PBF process).“Self-supporting angle varies depending on the material, but is usually around 45 degrees [8]” (ME process).“The de-facto value is 45 but most printers with some active cooling can handle a bit more. Try 55 and decrease if there is unwanted drooping, curling, or noodling’ [9]” (ME process).

Try 55 and decrease if there is unwanted drooping, curling, or noodling’ [9]” (ME process). “Angled surfaces (30~45): self supporting with rough surface finish [14]” (PBF process).The above variations of guidelines for overhangs are now re-written as structured rules, with antecedents and consequents:Overhangs (angular), if designed at greater than 45 degrees of undercut angle and built by a PBF process, are self-supporting.Overhangs (circular), if designed at proper range of tangent angles between 28 and 40 degrees corresponding to undercut radii and built by a PBF process, are self-supporting.

• Overhangs (hole), if modified with a peak under a proper range of undercut angles and built by a PBF process, are self-supporting.Overhangs (angular), if designed at less than around 0.075mm offset layering and built by a PBF process, are self-supporting.Overhangs (angular), if designed at greater than around 45 degrees of undercut angle and built by an ME process, are self-supporting.Overhangs (angular), if designed at greater than around 55 degrees of undercut angle, accompanied with unwanted drooping, curling, or noodling and built by an ME process, are self-supporting.Overhangs (angular), if designed at greater than around 30 degrees of undercut angle, aided by rough surface finish and built by a PBF process, are self-supporting.2.2 Modularizing Representation The rules presented in Section 2.1 can be re-written and represented using one single descriptive expression roughly formalized as:Category (type), if {conditions} then {consequences};where ‘category,’ ‘conditions,’ and ‘consequences’ include two components: primitives and modules.A primitive (or measured primitive) is a feature parameter such as ‘undercut angle,’ ‘undercut radius,’ ‘overhang distance,’ and ‘raw material type.’ Module is an implicit design feature such as ‘overhang,’ ‘support structure,’ and ‘surface finish.’ Accordingly, and more importantly, both primitives and modules are process-independent.

This means that design guidelines and rules that are written as prescriptive guidelines or explicit constraints can be re-established into design rules using only primitives and modules. In doing so, design rules can be interpreted more explicitly and implemented more efficiently because they are independent of context. Thereafter, they can be dynamically reconfigured, rather than created from scratch, from individual components for different AM builds and processes.Here, we propose a methodical approach to provide such modularity in design rules. Our approach is based on developing and introducing sets of fundamental building blocks. Such sets will be based on the following three premises:First, design rules should be established as process-independent sets of modular components.

Such sets will be based on the following three premises:First, design rules should be established as process-independent sets of modular components. For a specific AM build and process, these components can be specialized with process-specific parameters.Second, in process-specific implementations, design rules should be reconfigured and repurposed from existing discrete components, rather than created from scratch.Third, the principles abstracted from design rules should be robust enough to be impervious to fundamental changes or evolutions in layer-by-layer processes [23].A more specific explanation of these principles, and their use in a case study, will be provided in the following sections.3 THE METHODOLOGYDesign rules constitute a BoK that contains information useful for AM designers. In our view, the elements of design rules represent guidelines and constraints that make explicit how design and process information is managed, irrespective of the exact content.

In our view, the elements of design rules represent guidelines and constraints that make explicit how design and process information is managed, irrespective of the exact content. New design rules may be desired to effectively represent a new process, to incorporate new knowledge, or to develop in- house applications.Our methodology for design rule representation supports reusability, extendibility, readability, and computability, as shown in Fig. 1. As noted above, representations consist of well-defined components, consisting of two model classifications: primitives and modules. Primitives represent fundamental concepts of physical parameters observed in AM; modules establish dependencies between primitives and between primitives and other modules.

Primitives represent fundamental concepts of physical parameters observed in AM; modules establish dependencies between primitives and between primitives and other modules. Implementation-specific design rules can then be developed by• Defining the process-specific dependencies between modular components based on observed behaviorIdentifying the process-independent modular componentsAssigning values to modular components as appropriateFig. 1 Design rule principles and design rules with modularity for AMDesign rules are expressions and sentences composed of different language elements such as words, phrases, clauses, and sentences. Different languages and syntaxes can lead to variations in how design rules are constructed and interpreted. Design rules, written as prescriptive guidelines or explicit constraints specific to an AM-destined design [6-15], are more often than not dependent on a specific process.

Design rules, written as prescriptive guidelines or explicit constraints specific to an AM-destined design [6-15], are more often than not dependent on a specific process. A transformation to process-independency will result in generalized, modular expressions. This section describes both the principles and the formalisms needed to execute such a transformation.First, proper language elements in design rule expressions are re-defined into modular components and generalized as process-independent. The main purpose of design rules is to realizethe ‘design allowable.’ In general, a design rule is associated with two semantic notions: one is ‘designs (designed geometric features)’ implicitly generated or developed, and the other ‘guidelines or constraints’ explicitly defined or designated. For this reason, language elements in a design rule expression or representation fall, literally and semantically, into a category either of designs or of guidelines or constraints.

For this reason, language elements in a design rule expression or representation fall, literally and semantically, into a category either of designs or of guidelines or constraints. Thereafter they are classified into two components depending on the connotation they intend to carry; module conceiving implicit designs and primitive conceiving explicit guidelines or constraints. Table 1 shows some of the key attributes we considered in determining modules and primitives.Table 1. Types of attribute and their comparison of modular components moduleType of attribute implicit (designs) Semantic connotation Methods of generation derived or developed measured or defined Dependencyprimitive explicit (guidelines or constraints)yesnoSecond, design rules are re-written using conditional sentences, which provide descriptive expressions that form the basis for formalization.

They include both the ‘conditions,’ called hypothetical situations, and the ‘consequences.’ A full conditional sentence using a variety of grammatical forms and constructions generally contains two clauses: “the dependent clause expressing the conditions and the main clause expressing the consequences [24].” Here, an expression of design rules is formalized as a collection of ‘IF-THEN statements,’ as “the dependent clause is most commonly introduced by the conjunction ‘if’ [25].” It is written mathematically as is the main clause expressing the consequence. These two clauses, , consist of two modular components, primitives and modules, noted above.

These two clauses, , consist of two modular components, primitives and modules, noted above. This means that design guidelines and rules that are written as prescriptive guidelines or explicit constraints can be re-established into design rules with modularity using only primitives and modules.qthenpif;whereis the dependent clause expressing the conditions andqppandpqThe information needed to define primitives, modules, and their representations as conditional sentences are embedded in existing design rules, of which there are many [6-15]. That information is 1) embedded in the prescriptive guidelines or constraints contained in those rules and 2) represented with a number of different formalisms, including unstructured English. Currently, our approach relies on intuitively abstracting information from available guidelines and rules. Based on our interpretations, we then organize that information into primitives and modules – as defined below.

Based on our interpretations, we then organize that information into primitives and modules – as defined below. As we continue to develop our methodology, we will explore more deterministic methods for correlating primitives and modules.3.1 Primitives A primitive - independent in nature, measured or defined as designated, - is an explicit, single, identifiable entity. Primitives, however, are not necessarily interpreted in isolation since they are used to calculate a measure or designate a parameter in an AM process. Conceptually, primitives are “a set of standardized parts or independent units that can be used to construct a more complex structure [26].” This means that a primitive can exist with little or no customization at all. According to ASTM F2792, all seven categories of AM machines need a discretized version of the computer-aided design (CAD) geometry of a part that supports a layer-by-layer fabrication [5].

According to ASTM F2792, all seven categories of AM machines need a discretized version of the computer-aided design (CAD) geometry of a part that supports a layer-by-layer fabrication [5]. In addition, six among those seven need a discretized version of every single layer.qPrimitives can be classified into three groups: geometry, processes, and materials, depending on their source of generation. Several examples of each category are given in Table 2. These three categories of primitives are common to each of the seven ASTM F2792 categories of AM processes. As such, they are process-independent. They become process-dependent when specific values and ranges are acquired from direct observation of those processes. Both values and ranges must be verified before these primitives can be used to develop process-specific design rules.Table 2.

Both values and ranges must be verified before these primitives can be used to develop process-specific design rules.Table 2. Three different categories of primitives common to different AM processes Process o layer thickness o scan resolutions o build speed o build power o build orientations o platform dimensionsGeometry o feature dimensions o feature locations o feature orientations o undercut angleMaterial o raw material type o material density3.2 ModulesA module - or ‘context-to-primitive,’ derived or developed as dependent - is an implicit feature (often geometric), that is acquired from indirect observation (e.g., function) of given sets of either primitives or modules, or both. We have developed three types of modules.Type I: A module can be composed only from a set of primitives, for example:o A module for an overhang rule could be derived from two associated primitives,feature dimension and undercut angle, each of which has a certain range of values.

o A module for a surface finish rule could be derived from three associated primitives, layer thickness, scan resolutions, and raw material type, each of which has a certain range of values or types.o A module for a porous part rule could be derived from three associated primitives,layer thickness, scan resolutions, and raw material type, each of which has a certain range of values or types.Type II: A module can be composed from both primitives and other pre-developed modules,for example: o A module for a lattice structure rule could be developed from another pre-developed module, porous part, plus three associated primitives, feature dimension, feature locations, and feature orientations, each of which has a certain range of values.Type III: A module can be derived only from other pre-developed modules, for example:o A module for a chamfer rule could be developed from another pre-developed module,overhang.o A module for a support structure rule could be developed from two other pre-developed modules, overhang and surface roughness.Other examples that are common to each of those seven categories of AM processes, andthus process-independent, are also listed in Table 3.Table 3.

An example set of modules common to AM processesNo. of module Definition1 2 3 4 5 6 7 8 9surface finish overhang feature allowable shrinkage tolerance lattice structure support structure chamfer/fillet rib enforcementUsing both sets of modules and primitives, we can introduce ‘conditional sentences’ into design rules as written above. Conditional sentences will allow design rules to be tailored for specific implementations while maintaining the independent functionalities of modules. This clearly indicates that standardizing their fundamental principles will support customization for specific AM processes, technologies, and applications. The following case study provides an example of explicit interpretations of design rules specific to a PBF process.4 A CASE STUDY OF DESIGN RULE REDEFINITION WITH MODULARITY An example set of design rules, introduced from a commercial catalogue of design guidelines specific to a PBF process [14], is shown in Fig. 2.Fig.

2.Fig. 2 An example set of commercial design guidelines introduced for the case study [14]In this example, changes in the undercut angle of geometry impact the ability of the overhang to be self-supporting. If the undercut angle > 45 degrees, the overhang will prove self-supporting. If the undercut angle falls in an interval (30 degrees to 45 degrees), then the overhang will only be self-supporting provided it has a rough surface finish. The subsidiary physical interrelations between the surface features (roughness) and the surrounding powder material makes self- supporting possible.4.1 Primitive Definitions An example set of primitivesPmade for the case study is given in Table 4, where elementsPbelonging to Short definitions of the primitive are also given in Table 4., among others, are defined as process-independent over different AM processes.Table 4.

An example set of primitivesPdefined for the case studyElement symbol1PAbbreviation UC(undercut angle)DefinitionAn angle created by offset layering build2PSR(scan resolution)A process resolution on x-y plane3PRM(raw material type) A characteristic of the material used4PFD(feature dimension) A dimensional scale of a feature during build5PLT(layer thickness)A layer thickness during build4.2 Module Definitions An example set of modulesMmade for the case study is given in Table 5. Elements belongingMto f(x) is function of x., among others, are derived fromPand other pre-derived modules belonging to {M}, andTable 5. An example set of modulesMderived for the case studyElement symbol1MAbbreviationOH(overhang)Function definition ,1P4Pf ()2MSF(surface finish)f (2P,3P5P)3MSS(support structure)f (1M,2M)The process of generating an example set of modulesMis illustrated in Fig.

3. A module1M(OH: overhang) can be derived from two associated primitives,1P(UC: undercut angle) and4P(FD: feature dimension), each of which has a certain range of values. Another module2M(SF:surface finish) can be derived from three associated primitives,2P(SR: scan resolution),3P(RM:raw material type), and5P(LT: layer thickness), each of which has a certain range of values ortypes. Finally, the third module3M(SS: support structure) can be derived from1Mand2Mthatare previously derived. Elements ofMare process-independent over different AM processes.Fig. 3 A set of design rules studyRgenerated using modular componentsPandMfor the case4.3 Redefinition of Design Rules An example set of design rules,R, redefined from the ones written above and shown in Fig.Rare derived from both 2, is given in Table 6.

Elements belonging to relationships between the dependencies are outlined by the design principles.PandR, whereRTable 6.

An example set of design rules Element symbolscompared with different representationsType of representation Prescriptive guidelineDesign rule representation1RIf given an Overhang with an Undercut of less than 30 degrees, then Support Structures are needed.IF-THEN statementIF {1Mwith1P} THEN {3M}Description logic1R:1M(OH1)3MLessThan (UC1,30)  (SS1)2RPrescriptive guidelineIf given an Overhang with an Undercut of greater than 30 degrees and less than 45 degrees and the surface roughness is less than ‘1’, then Support Structures are needed.IF-THEN statementIF {1Mwith1P2Mwith3P} THEN {3M}Description logic2R:1M(OH1)GreaterThan (UC1,30)LessThan (UC1,45)2M(RS1)3MLessThan (RS1,1)  (SS1)The consequents are ‘functions of’ the antecedents, and the exact relationship established will depend on the process and the intent of the rule.5 CONCLUSIONS This paper proposed a methodical approach to developing design rules with modularity for additive manufacturing.

We demonstrated that the use of process-independent representations provide a basis for explicit interpretations of specific design rules, specifically for the powder-bed fusion process. Conclusions about design rules with modularity have been made as follows:  First, they enable repeatability in an AM design guideline under different and diversedesign and process environments.Second, they can communicate process-independent design principles to users that are unfamiliar with AM processes.Third, they are comprehensible and, therefore, manageable when subjected to changes and updates of AM machines and processes.Modularity is applied to not only the geometric representations of the design, but also to the data providing support for the manufacture of the design. By defining design rules as sets of modular components, they can be more easily interpreted and implemented, independent of context.

By defining design rules as sets of modular components, they can be more easily interpreted and implemented, independent of context. Since design rules impact process planning as well, we will also look to abstract recurring themes out of the process-specific guidelines.Future research will continue investigating the use of our methodology on more complicated examples. We will demonstrate the reconfiguration of design rules from individual components, including process-specific parameters for different AM builds and processes. We will promote the consistent application of design rule principles as a means for mitigating ambiguities that may be introduced both within and between processes. We will continue to develop an extended design automation paradigm for the design and fabrication of new AM parts.REFERENCES [1] http://www.merriam-webster.com/dictionary/rule (accessed June 30, 2015). [2] Gausemeier, J., Echterhoff, N., Kokoschka, M., Wall, M.

SIGN SYNTHESIS OF ADAPTIVE MESOSCOPIC CELLULAR STRUCTURES WITH UNIT TRUSS APPROACH AND PARTICLE SWARM OPTIMIZATION ALGORITHMHongqing Vincent Wang a R&D EngineerChris Williams b Research AssistantDavid W. Rosen b* Professora IronCAD Inc., 700 Galleria Parkway, Suite 300, Atlanta, GA 30339 b The George W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology Atlanta, GA 30332Corresponding Author: 404-894-9668, david.rosen@me.gatech.edu.ABSTRACTCellular material structures have been engineered at the mesoscopic scale for high performance and multifunctional capabilities. However, the design of adaptive cellular structures - structures with cellular configurations, sizes, and shapes designed for a specific geometric and loading context - has not been sufficiently investigated. In this paper, the authors present a design synthesis method with the use of unit truss approach and particle swarm optimization algorithm to design adaptive cellular structures.

In this paper, the authors present a design synthesis method with the use of unit truss approach and particle swarm optimization algorithm to design adaptive cellular structures. A critical review is presented to show the pros and cons of the new design synthesis method and an existing homogenization method. The research extends the application of additive manufacturing in the design of new materials for high performances and benefits its long-term growth.KEY WORDSMesoscopic Cellular Structure, Adaptive Cellular Structure, Design Synthesis, Unit Truss,Particle Swarm Optimization, Homogenization Method, Additive Manufacturing1 CELLULAR MATERIALS WITH DESIGNED MESOSTRUCTUREMore than 50 years of research have been dedicated to the development of design and manufacturing methods that are capable of applying the advantages of cellular materials found in the natural world to engineered artifacts.

Much like wood, bone, and coral, man-made materials with cellular mesostructure (cells on the order of 0.1 to 10mm) such as foams, honeycombs, and lattices feature gaseous voids that are thoroughly dispersed throughout the solid phase of its body [1]. The key advantage offered by cellular materials is their high strength and relatively low mass. The concept of artificial cellular materials with designed mesostructure is motivated by the desire to put material only where it is needed for a specific application. An increase in strength and decrease in relative density are laudable benefits of this motivation; however, mesostructured materials are promising because they can be designed and manufactured for specific, multiple design objectives and product functions.Due to the limitations found in traditional manufacturing technology, most available engineered lightweight structural cellular materials are linear patterns of primitives.

With the advance of additive manufacturing technologies, however, engineers have the freedom to design structural cellular materials to provide stiffness, energy absorption characteristics, and good thermal and acoustic433insulation properties by orienting strut directions and adjusting strut diameter and cross-section (Figure 1a) [2]. Seepersad and coauthors, for example, designed the topology of an extruded cellular material to find the best compromise between heat transfer and part strength in a structural heat exchanger application [3]. Williams and coauthors investigated additive manufacturing processes and selected Selective Laser Melting and Three-Dimensional Printing for the manufacture of parts with metallic designed mesostructure [4].Compliant mechanisms, another type of cellular structure (Figure 1b), can be designed to transform motions, forces, or energy by elastic deformation as opposed to traditional mechanisms consisting of rigid links [5-7].

Living organisms such as cartilage and living cells are sophisticated compliant mechanisms and they possess both a high degree of moving freedom and the ability to manipulate that freedom [8, 9]. Compliant mechanisms in nature are able to carry out diverse and soft movements. Most engineered compliant mechanisms are limited due to inferior design methods for complicated mechanisms that do not simultaneously consider axial forces, bending, torsion, nonlinearity, or buckling. Most compliant mechanisms can only carry out relatively simple movement compared to living organisms.a. A lightweight truss structure supporting car body for high strength and minimum deflectionb. A compliant mechanism as a new material structure with negative Poisson’s ratioFigure 1.

A lightweight truss structure supporting car body for high strength and minimum deflectionb. A compliant mechanism as a new material structure with negative Poisson’s ratioFigure 1. Examples of cellular structuresCellular structures are called “adaptive” if their configuration, geometry, and sizes are chosen to significantly enhance the structures’ performance, while conforming to the shape of parts they are embedded within or of neighboring parts. The systematic design of adaptive cellular structures has not been studied extensively. In this paper the authors propose a method for the design synthesis of adaptive cellular structures. Specifically, a synthesis method using a unit truss approach and integrated with particle swarm optimization is developed to systematically design adaptive mesoscopic cellular structures. The core components of the design method, the unit truss approach and particle swarm optimization method, are presented in Sections 2 and 3, respectively.

The core components of the design method, the unit truss approach and particle swarm optimization method, are presented in Sections 2 and 3, respectively. The integration of these core constructs is described in Section 4. In Section 5, the method is validated through its application to a cantilever beam example and its comparison against the homogenization method with optimality criteria. Finally, in Section 6, the methodology is tested against a more complex example inspired by a morphing wing technology. The closure is offered in Section 7.2 UNIT TRUSS APPROACHA few structural analysis approaches have been developed to analyze compliant mechanisms. Typical methods include the ground truss (discrete) approach and the homogenization (continuum)434method. The ground truss approach can only provide a rough estimate for the geometry of designed structures [10-12].

The ground truss approach can only provide a rough estimate for the geometry of designed structures [10-12]. The homogenization method achieves better results by using artificial unit cells; however, it may result in non-realizable elements and can be computationally expensive [13, 14].We propose the unit truss, a new type of unit cell, to analyze cellular structures. The unit truss consists of the central node and a set of the half-struts that are connected to that node. An example of a unit truss is shown in Figure 2 [15]. This new unit cell approach was developed to accurately and efficiently analyze lightweight structures and compliant mechanisms, and to support their systematic design. The unit truss is leveraged from the ground truss approach and homogenization method. A microstructure cell is used to represent the material distribution in the homogenization method. An advantage of the unit truss is that it can be used as both the cell primitive for analysis and synthesis.

An advantage of the unit truss is that it can be used as both the cell primitive for analysis and synthesis. Furthermore, the unit cell is manufacturable, whereas the microstructures in the homogenization method are artificial and not manufacturable [16].Unit Truss 1Unit Truss 2Unit Truss 3Figure 2. Definition of unit trussIn order to analyze conformal cellular structures, the mechanics model of unit trusses was successfully developed. As shown in the stress plot of a sample unit truss in Figure 3, the strain and stress around the nodes are complicated due to considerable inter-strut interactions and large bending moments [15]. As such, the unit truss model includes considerations of axial forces, bending, torsion, nonlinearity, and buckling [15].(2)1u(1)1u(1)3u(2)2u(2)3u(1) 2u(0)1u(0) 2u(3) 2u(3) 1u(5)3u(0)3u(3)3u(5)1u(5)2u(4)3u(4)1u(4) 2uFigure 3.

Using unit trussThe constitutive equations of 2-D and 3-D unit trusses are derived using beam theory as shown represents the nodal in Equations 1-3. displacements and forces. Each strut is modeled as a conventional frame element from finite element analysis methods; however, we have added provisions for modeling geometric and materialdenotes the linear elasticity of a unit truss, while U (cid:4)and F (cid:4)eK (cid:4)435nonlinearities, as well as geometric interactions at nodes. Unit trusses can have any number of incident struts. They can be thought of as special elements with which to analyze large cellular structures using methods similar to conventional finite element methods.

Material nonlinearity occurs due to nonlinear stress-strain behavior. Geometric and material nonlinearities are considered with the tangent stiffness method. The tangent stiffness method is a linearization approach to analyze the nonlinear behavior of compliant mechanisms. Linear elasticity theory is used to solve nonlinear problems. The behavior of an elastic unit truss can be traced back incrementally using Equation Equation 4, which is in a linear form [17]. are the tK (cid:4) incremental nodal forces. Both geometric and material elastic nonlinearities are considered using represent Equation Equation 5.

Compared to the homogenization method, the unit truss approach is more efficient due to fewer microstructures used for analysis. As a result of this research, the analysis of lightweight structures and compliant mechanisms can utilize the same mechanics model (unit truss). This enables us to design structures for both stiffness and compliance simultaneously.3 PARTICLE SWARM OPTIMIZATIONThe design synthesis of adaptive cellular structures is a large-scale nonlinear problem with multiple objectives and a mixed-discrete design space. The design problem formulations for lightweight structures and compliant mechanisms are shown in Figure 4. Strut sizes (diameter for cylindrical struts or width for 2-D rectangular struts) are the design variables represented by436eK (cid:4)., where n is the number of struts in the starting structure topology. The constraints ix i ( include the bounds on strut diameters, static equilibrium, and stress.

The constraints ix i ( include the bounds on strut diameters, static equilibrium, and stress. The objectives are to minimize the normalized Mean Squared Deviation (mean( between the desired shape and the actual SD shape under deformation, and to minimize the normalized total material volume Vnorm [18]. Detailed explanation of this design problem is given in the next section.n 0,1,..., )(cid:32)))knormParticle Swarm Optimization (PSO) and Genetic Algorithms (GA) were selected from available optimization algorithms to systematically search for design solutions. PSO shares similarities with GA, while PSO enables cooperative behavior among individual design instances as well as the competition modeled using GA.

Lightweight structuresb. Compliant mechanismFigure 4. Problem formulation for cellular structuresConceptually, PSO simulates the movement of birds in a flock, where individuals adjust their flying according to their experience and other individuals’ experiences during searches for food [20]. It starts with global search and then gradually shifts to local search through the change of weight coefficient. Each of the p particles (birds), which represent design instances, is given an initial position and velocity. In the context of this design problem, an initial position corresponds to a set of strut diameter values for the entire mechanism and the velocity is the change rate of the strut diameters between design iterations. The key step in the PSO algorithm is the update of the velocity from one iteration to the next.

Positions are computed simply using Equation 7. During each iteration, all particles’ velocities and positions are updated, and then the objective function is computed. Convergence is checked by considering changes in objective function values, as well as design variable changes. The authors developed the design synthesis method by integrating PSO with the unit truss approach to systematically design compliant mechanisms with multiple inputs/outputs. In the following section, the implementation of this method is described.x x v = + i i iEquation 74IMPLEMENTATIONThe design synthesis was implemented and integrated in MATLAB. The PSO algorithm best x represents the ( ) kf represents the )kfollows the general structure shown in Figure 5. “DV” means design variables.

Usually,is set as zero. The maximum number of iterN to avoid a dead loop. Automatic termination of the search iterations for each run is given as process is applied when the target is reached as given in Equation 8 or the convergence deviation (such as standard deviation of the search solutions) is less than the tolerance of design variables var(cid:72) as given in Equation 9.fgoalobj(cid:72) is the tolerance of objective function, andvar(cid:72) is the tolerance of design variables.fg(cid:16)fgoal(cid:100)(cid:72)objEquation 8mean v (cid:72)(cid:100) ()varEquation 9startw and the endw are given based on experience. The recommended values are given as startw and endw will make the particles jump with bigger steps and have 2(cid:77) balance 2(cid:77) are integer 2.startw to endw . The start velocity inertiaThe velocity inertiakw linearly varies fromend velocity inertiastartw(cid:32)0.9toendw (cid:32).

Implementation of PSO for adaptive cellular structure designDN is the number of design variables, in another words, the number of struts. The number of PN PN , is given based on experience. Usually it is given as 20. A larger number of particlesparticles, can make the search result more reliable, but this also requires more iterations to converge.439,The maximum moving velocity maxv and lower bound of design variables. In fact, maxv particles.is usually set as the difference between the upper boundcan be looked as the largest allowable jump step for5 VALIDATIONIn order to validate the integration of the unit truss approach and particle swarm, we choose to compare the results of its implementation on a simple problem with those from the Optimality Criteria (OC) with homogenization method [21]. The validation example is shown as in Figure 6, which is a 300x150x5 mm cantilever plate with elastic modulus of 200MPa.

The validation example is shown as in Figure 6, which is a 300x150x5 mm cantilever plate with elastic modulus of 200MPa . It is loaded with a 1 N force at a distance 40 mm from the top edge. The goal is to optimize the material distribution with a material usage of 30% of the domain volume for minimum deflection.1N150mm300mmFigure 6. Validation example for design synthesisThe solution generated by the Optimality Criteria and homogenization method for a desired average density of is shown in Figure 7. In the homogenization problem, each element has one design variable, the density. Elements are 10x10 mm in size, thus the plate has 450 elements and variables. The 99-line code of OC with homogenization method by Sigmund [37] is used to solve the problem for the comparison. Convergence is achieved within 80 iterations - about 4 minutes elapsed time on a standard PC with a 2.4GHz processor.

Convergence is achieved within 80 iterations - about 4 minutes elapsed time on a standard PC with a 2.4GHz processor. The maximum deflection found using this solution strategy was 1.9 mm.= 0.3(cid:85)The solution of the same problem using our PSO method with unit truss approach for the * is presented in Figure 8. This structure has the same thickness (5mm) as average density that for homogenization method. The unit truss problem has 58 struts, wherein the strut width is the sole variable for each. The PSO with unit truss approach is run for 200 iterations with 20 particles and the resulting deflections ranged between 1.3 mm and 1.8 mm. In each iteration, the objective function is evaluated 20 times since 20 particles are used. For 200 iterations and 20 particles, the objective function is evaluated 4000 times, and about 12 minutes elapsed time was required to solve the problem.

For 200 iterations and 20 particles, the objective function is evaluated 4000 times, and about 12 minutes elapsed time was required to solve the problem. The thin lines shown in Figure 8 represent the struts with negligible width.= 0.289(cid:85)The constraint(cid:85) (cid:100)0.3is applied in the search process by using penalty function and considered as a continuousfunction; hence, the search result may not have exactly(cid:85) (cid:32)0.3.440Figure 7. Result from homogenization methodFigure 8. Result from the new design synthesis approachThe unit truss approach returned a solution that had a slightly smaller area and deflection than that of the homogenization method. Though the unit truss problem had an order of magnitude fewer variables, it took 3 times longer to reach a solution.

Though the unit truss problem had an order of magnitude fewer variables, it took 3 times longer to reach a solution. The PSO with unit truss approach took many more runs (4000 runs; 200 iterations with 20 particles) to evaluate the objective functions than the OC with homogenization method (80 runs). The main reason for this is that the OC uses gradient and Hessian matrices as the heuristic information during the search process, while the PSO does not use either of them. Regarding the geometry of the solutions, the result with the unit cell approach is manufacturable, while the gray elements in the result from homogenization are not manufacturable. This is a limitation of the homogenization method.Both the homogenization method and the unit truss approach have their advantages and disadvantages; however, the homogenization method with optimality criteria is very effective for continuous convex design domain.

It needs gradients and Hessian matrices, which are obtained only with expansive efforts and computational resources. The homogenization method also presents results that are not manufacturable. For the validation example above, the design domain is continuous with design variables varying from 0.01 to 1.0, but those elements with design variables ranging between 0.01 and 1.0 are not manufacturable or do not physically exist. To precisely represent the physical meaning, the design variables are not continuous and the design domain is discrete. Moreover, the gradients and Hessian matrices used for homogenization method are obtained by an approximation with power-law interpolation [21]. This power-law interpolation cannot precisely represent the true physical behavior of microstructures. When solving 3-D structural problems, the number of design variables for homogenization method is scaled much more dramatically than unit truss approach.

When solving 3-D structural problems, the number of design variables for homogenization method is scaled much more dramatically than unit truss approach. For example, 8000 microstructures (lateral size 10mm) are needed to represent a 3-D domain with lateral size 200mm. As a result, 8000 design variables are required. The research about 3-D structural homogenization method has not been explored as extensively as 2-D homogenization method because of those difficulties.The unit truss approach with particle swarm optimization can solve problems with discrete design domains. It does not require the gradients and Hessian matrices, which makes it a convenient method for users. Furthermore, the number of design variables in the unit truss approach is much fewer than the homogenization method. The number of design variables is only about 11% of that of homogenization method for the validation problem.

The number of design variables is only about 11% of that of homogenization method for the validation problem. The search time of the design synthesis process should be cut significantly with the use of unit truss approach.4416 EXAMPLE PROBLEM: MORPHING WINGFor further validation of our methodology, we investigated the design synthesis of a compliant mechanism that enables an entire closed-loop airfoil profile to change shape from NACA 23015 to FX60-126. The nonlinear behavior of largely compliant mechanisms under large deformation is considered. The resulting design is validated by testing its robustness and considering nonlinearity.A morphing wing concept for AAI's Shadow shown in Figure 9 is proposed as an example problem in this research. AAI's Shadow is a small Unmanned Aerial Vehicle (UAV) for information collection [22]. The flight range and endurance of UAV are limited by the fuel storage capacity.

AAI's Shadow is a small Unmanned Aerial Vehicle (UAV) for information collection [22]. The flight range and endurance of UAV are limited by the fuel storage capacity. It is greatly desired to increase the flight range and endurance without the addition of fuel. As the fuel is burned throughout the mission, the total weight of the UAV decreases. As the aircraft’s weight decreases, the wings’ working condition changes, and a different airfoil shape would better serve the aircraft. The airfoil geometry is desired to morph and accommodate the changing working condition for high airfoil performance. Wings with adaptive shapes can minimize the mission drag and improve the fuel efficiency. In the AAI’s Shadow example studied by Gano and Renaud [23], the wing cross- section morphs from NACA 23015 to FX60-126 are as shown in Figure 11 [24]. NACA 23015, represented by the large profile, is bulky and has more capacity to store fuel as the starting cross- section.

NACA 23015, represented by the large profile, is bulky and has more capacity to store fuel as the starting cross- section. FX60-126, represented by the solid block, is slender as desired for the shape at the end of mission.

AAI's shadow 400 UAV system [26]Figure 10. Problem formulation of compliant mechanism for morphing wingThe problem formulation of the design synthesis for the morphing wing is shown in Figure 10 ) of the struts in the initial topology that has [24]. The design variables are the widths a total of 117 struts and72 nodes. Both the most left and right nodes are fixed; five pairs of equal and opposite forces are applied on the upper and lower boundary as shown in Figure 12. The design objective is to minimize the weighted sum objective value f x of the weighted mean squared ( ) deviation (mean( . The displacements are sampled from 8 points ) on the airfoil shown in Figure 11.

Each evaluation of the objective function takes about 0.612 seconds. The experiment computer has an Intel P4 2.4GHz CPU and 512MB RAM. The best run produced the objective function value f(x) = 4.37 with the deflections at the sampled points shown in Table-1 and the resulting deformed shape is shown in Figure 13.Table-1. Deviations between actual and target deflections at the sampled points 16Node ID Target12 6.580 11.400 13.700824 1.56028 2.14033 0.00063 9.43069 -3.6506.397 -11.360 -13.759 -1.686 synthesized Deviation -12.977 -22.760 -27.459 -3.246 1.250 -0.311 AfterActualBefore2.250 -4.390 3.125 0.9850.020 -9.427 3.645 -0.020 -18.857 7.295 6.848 3.031 0.225 -2.582 3.875 3.031A topology cleaning process is performed on the resulting synthesized topology by removing “zero-width” struts, defined as widths close to 0.00025mm. During the cleaning process, the dangling elements are removed as well.

During the cleaning process, the dangling elements are removed as well. The cleaned structure has 68 struts and 60 nodes as shown in Figure 14. After cleaning, the objective function value f(x) = 5.012 is a little larger than that before cleaning. This difference is caused by the removal of struts that contribute to the structure’s stiffness even if their widths are very small. It is important to note, however, that this does not extensively influence our design synthesis result.443f(x) = 4.367, mean(SDk) = 2.854, Vnorm = 0.302 Figure 13. Deformed structure with five pairs of opposite forces after synthesisf(x) = 5.012, mean(SDk) = 3.69, Vnorm = 0.264 Figure 14. Cleaning topology by removing zero-width struts7 CONCLUSIONSThe authors successfully developed a design synthesis method to design adaptive mesoscopic cellular structures with the use of unit truss approach and Particle Swarm Optimization algorithm (PSO).

The new design synthesis method enables the analysis and design of complicated cellular structures including lightweight structure and compliant mechanisms. Axial forces, bending, torsion, nonlinearity, and buckling can be simultaneously considered. The design synthesis of adaptive cellular structures is a large-scale nonlinear problem with multiple objectives and a mixed-discrete design space. The authors formulated the design problems for both types of adaptive cellular structures, and selected PSO to systematically search for design solutions. The unit truss approach and PSO were successfully implemented and integrated in MATLAB.The authors solved a cantilever beam problem by using both the new design synthesis method and the existing homogenization method.

The resulting designs are similar in terms of average density and deflection; however, all struts in the result of the design synthesis method with the unit cell approach are manufacturable, while some of the elements in the result created with the homogenization method are not. As opposed to the homogenization method, the new design synthesis method does not need efforts to obtain gradients or Hessian matrices, and has fewer design variables. The authors demonstrated the new design synthesis method through the successful design of a compliant mechanism for morphing wing, which has multiple objectives and a nonlinear design domain.ACKNOWLEDGEMENTSWe gratefully acknowledge the U.S. National Science Foundation, through grants IIS-0120663 and DMI-0522382, and the support of the Georgia Tech Rapid Prototyping and Manufacturing Institute member companies for sponsoring this work.REFERENCES[1] J. Banhart, "Manufacturing Routes for Metallic Foams," Journal of the Minerals, vol.

Banhart, "Manufacturing Routes for Metallic Foams," Journal of the Minerals, vol. 52, pp. 22-27, 2000.[2] H. Wang, Y. Chen, and D. W. Rosen, "A Hybrid Geometric Modeling Method For Large Scale Conformal Cellular Structures," ASME Computers and Information in Engineering Conference, Long Beach, California, September 24-28, 2005, 2005, Paper Number: DETC02/CIE-34495[3] C. C. Seepersad, "A Robust Topological Preliminary Design Exploration Method With Materials Design Applications," PhD, School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA, 2004.[4] C. B. Williams, F. Mistree, and D. W. Rosen, "Investigation of Solid Freeform Fabrication Processes for the Manufacture of Parts with Designed Mesostructure," ASME IDETC Design for Manufacturing and the Life Cycle Conference, Long Beach, California, 2005[5] S. Kota, J. Joo, Z. Li, S. M. Rodgers, and J.

K. Rule, "Automatic truss design by optimized growth," Journal of Structural Engineering,vol. 120, pp. 3063-3070, 1994.[12] A. G. M. Michell, "The limits of economy in frame structures," in Philosophical MagazineSect. 6, vol. 8, 1904, pp. 589-597.[13] M. P. Bendsoe, Optimization of structural topology, shape, and material, Springer-Verlag,Berlin Heidelberg, 1995.[14] M. P. Bendsoe and N. Kikuchi, "Generating optimal topologies in structural design using a homogenization method," Computer Methods in Applied Mechanics and Engieering, vol. 71, pp. 197-224, 1988.[15] H. Wang, "A unit cell approach for lightweight structure and compliant mechanism," PhD Dissertation, School of Mechanical Engineering, Georgia Institute Of Technology, Atlanta, GA, 2005.[16] G. I. Rozvany, Topology Optimization in Structural Mechanics: Springer Verlag, 1997. [17] W. McGuire, R. H. Gallagher, and R. D. Ziemian, Matrix structural analysis, 2nd ed, JohnWiley, 2000.[18]J. Neter, M. H.

21, pp. 120-127, 2001.[22]"Materials for lean weight vehicles: the battle for a better body," Steel Times (UK), vol. 226,pp. 73-74, 1998.[23] S. E. Gano and J. E. Renaud, "Optimized Unmanned Aerial Vehicle With Wing Morphing For Extended Range And Endurance," 9th AIAA/ISSMO Symposium and Exhibit on Multidisciplinary Analysis and Optimization, Atlanta, GA, September 4-6, 2002, 2002, AIAA-2002-5668[24] H. Wang and D. W. Rosen, "An Automated Design Synthesis Method for Compliant Mechanisms with Application to Morphing Wings," ASME Design Mechanisms and Robotics Conference, Philadelphia, PA, September 10-13, 2006, 2006, DETC2006-99661[25] M. Selig, "UIUC Airfoil Coordinates Database - Version 2.0," Department of Aerospace Illinois 61801Engineering, University of http://www.ae.uiuc.edu/m-selig/ads.html, August 10, 2005, 2005Illinoisat Urbana-Champaign, Urbana,[26]"Shadow 400 UAV system," AAI Corporation http://www.aaicorp.com, September 13, 2005445.

Lane1, S. Mekhontsev2, S. Grantham2, M. L. Vlasea3, J. Whiting1, H. Yeung1, J. Fox1, C. Zarobila2, J. Neira2, M. McGlauflin1, L. Hanssen2, S. Moylan1, A. Donmez1, J. Rice21Engineering Laboratory, 2 Physical Measurement Laboratory, National Institute of Standards and Technology, Gaithersburg, MD 208993Mechanical and Mechatronics Engineering, University of Waterloo, Waterloo, ON, Canada N2L 3GIAbstractThe National Institute of Standards and Technology (NIST) is developing a facility titled the Additive Manufacturing Metrology Testbed that will enable advanced research into monitoring, controls, process development, and temperature measurement for laser powder bed fusion additive manufacturing and similar processes. This system provides an open control architecture as well as a plethora of sensor systems and calibration sources that are primarily radiance-based and aligned co-axially with the laser beam and focused on the laser interaction zone.

This paper briefly reviews the system requirements, and details the current progress of the facility design and construction. Mechanical, optical, and control systems designs are detailed with select highlights that may be relevant to additive manufacturing researchers and system developers. Recent experimental results from the prototype laser control and in-situ monitoring system are also highlighted.IntroductionVlasea et. al. described the proposed construction of an additive manufacturing metrology testbed at the National Institute of Standards and Technology (NIST) that would provide a fully controlled research-oriented laser powder bed fusion (LPBF) system to enhance research capabilities in additive manufacturing (AM) and materials research [1].

This system is primarily designed to serve two different but supporting research endeavors by NIST’s Engineering Laboratory (EL) and Physical Measurement Laboratory (PML), which require well-controlled and characterized high laser energy input applied to solid metals, metal powders, or other materials, and simultaneous measurement from a plethora of primarily optically-based metrological instruments. For EL’s use, the system is given the moniker Additive Manufacturing Metrology Testbed (AMMT), and for PML’s use, the system is named Temperature and Emittance of Melts, Powders, and Solids (TEMPS).Certain commercial entities, equipment, or materials may be identified in this document in order to describe an experimental procedure or concept adequately.

Such identification is not intended to imply recommendation or endorsement by the National Institute of Standards and Technology, nor is it intended to imply that the entities, materials, or equipment are necessarily the best available for the purpose1145Overarching goals for the system are as follows:1. Build an open-platform LPBF testbed instrumented for fundamental process study and optimization, as well as optical metrology (AMMT).2. Research the measurement science of various real time monitoring and control methods for measuring and improving AM build quality (AMMT).3. Establish new emittance measurement capability for materials in various stages of aggregation over solid and liquid-phase temperatures (TEMPS).4. Develop new NIST calibration services and standard reference data on polymers, ceramics, high temperature alloys, composites, and coatings (TEMPS).5.

Develop new NIST calibration services and standard reference data on polymers, ceramics, high temperature alloys, composites, and coatings (TEMPS).5. Develop capacity to accurately measure temperature distribution across the melt pool under laboratory conditions (common goal).6. Study methods for real-time process thermometry in the production environment (common goal).Since the previous report [1], the concept has progressed from definition of goals and conceptual design, to near final mechanical and optical design, and initialization of the research facilities construction, system control architecture, and mechanical fabrication and assembly. In addition, initial controller programming and laser scanning tests on a prototyping system have been performed. The focus of this paper is to provide a general update on this progress. A comprehensive description of all design aspects would be too lengthy, so an ‘overall’ system design concept is provided here.

A comprehensive description of all design aspects would be too lengthy, so an ‘overall’ system design concept is provided here. Highlights are provided regarding ideas and results that may be pertinent to the additive manufacturing research community. While the TEMPS-based operations and design concepts will support additive manufacturing research, the focus of this paper is primarily on aspects pertaining to AMMT operation.Design FunctionalityVlasea et al. described the general dual-functionality of AMMT/TEMPS system as operating in either build mode or radiance-based metrology mode [1]. In build mode, which is the primary mode for AMMT operation, the system is a fully controllable LPBF machine, with fully definable and controllable build parameters including but not limited to laser power, scan speed, layer thickness, and scan strategy.

Primary focus of AMMT metrological research will be on radiometric measurements using optical instruments in a co-axial configuration, that is, aligned with the laser and monitoring dynamically changing radiant emission from the melt pool. Design flexibility also allows for staring configurations (as opposed to co-axial) for imaging or other radiometric sensors, which are either mounted within the vacuum chamber or externally mounted looking through viewports. These two configurations, allowing for melt pool monitoring or layer- wise imaging, are already being incorporated into commercial LPBF systems [2]. The advantage brought by the AMMT, apart from full process control, characterization, and reconfiguration flexibility, is that radiometric calibration sources will be located adjacent to the build bed, allowing for rapid in-situ calibration.

In addition, processes and instrumentation composing the TEMPS systems will provide additional references and comparison.1146In radiance based metrology mode, which is the mode for TEMPS-based operation, the system is reconfigured such that other optical beam lines and instrumentation are accessed. Similar to the co-axial configured instruments used in AMMT operation, the TEMPS-based operation will utilize emitted radiation from the heat affected zone transmitted to co-axially aligned optics and instruments.

However, unlike in AMMT operation, which utilizes visible and near-infrared spectrum, TEMPS operation is broad band (visible to long-wave infrared), and thus requires special reflective or broadband optics [3].This operation mode is intended to extend current capabilities at NIST for material optical property measurement [4], and allow measurement at higher temperatures, operation in vacuum or high purity inert environment, measurement at shorter wavelengths, and measurement of material phases other than solids. These extended phase measurements primarily focus on powders and liquids, but availability of high laser energy densities may allow for vapor and plasma characterization as well. One of the measurement procedures, titled the DYnamic Meltpool Emissometry (DYME) method, essentially replicates procedures used at NIST for spectral emissivity characterization of solid material samples [5].

However, the DYME method utilizes the moving melt pool (or non-liquid, laser heated spot) as a target, statically held in the field of view of the co-axial optical system, rather than a statically heated sample. This method, compared to previous methods using larger, isothermally heated samples, reduces oxidation, contamination, and/or evaporation due to the small area of the heat affected zone. These measurements have direct application to LPBF process monitoring and research, and details of the TEMPS-specific system design, operation, and measurement procedures will be provided in separate publications.System Modules and Mechanical DesignAlthough complex, the overall system allows expanded flexibility for researchers to reconfigure subsystems individually by providing a modular and separable design. Figure 1 shows an external view of the computer aided design (CAD) assembly, in which the carriage is pulled out of the vacuum chamber onto the transfer dolly.

Figure 1 shows an external view of the computer aided design (CAD) assembly, in which the carriage is pulled out of the vacuum chamber onto the transfer dolly. The figure highlights the five main system modules. The removable carriage allows researchers to have better access to assemble, reconfigure, and clean subsystems on the carriage externally from the vacuum chamber.1147Optics Tower (cover removed)Process Control Module (cover removed)Carriage Dolly Main Carriage Vacuum ChamberBeam path to TEMPS Module(not shown)Figure 1: Solid model assembly of the AMMT/TEMPS system with the main carriage removed from the vacuum chamber. Not shown are the TEMPS module, which is positioned behind the vacuum chamber, and various cables and plumbing.The main carriage is the main mechanical system built around a stainless steel frame and granite base, and houses the build bed, powder feed, recoating mechanisms, and most major mechanical and motion components.

The carriage frame can be removed from the vacuum chamber by rolling along caster wheels on vee and flat tracks, and onto the external, mobile carriage dolly, which has similar tracks. The carriage dolly includes a winch and strap that connects to the main carriage to pull the carriage out of the vacuum chamber.The vacuum chamber, designed for medium vacuum (0.1 Pa or 10-3 torr) operation, provides vacuum purging prior to backfilling with nitrogen or argon inert gas. This provides faster means to reduce oxygen levels compared to simply purging with inert gas. The vacuum chamber also provides access panels on all sides for cable and hose feedthroughs and user access to the build area and carriage underside.

The vacuum chamber also provides access panels on all sides for cable and hose feedthroughs and user access to the build area and carriage underside. This also provides flexibility for incorporating future custom panels with ports or feedthroughs for newer or varying instrumentation or research projects.The optics tower contains the main optical components to control and guide the heating laser, and divert back-emitted melt pool radiation to different modules. Pulling vacuum on the chamber will cause it to deflect, therefore the optics tower and fixed optical components are mechanically de-coupled from the chamber, and coupled to the granite by sitting on the optics tower bridge (not shown in Figure 1). The optics tower and internal components are vacuum compatible and can be purged, however a broadband window can be inserted between the optics1148tower and main chamber to isolate the optics tower from vacuum.

The optics tower contains the galvanometric (galvo) scanner mirrors, removable beamsplitter/reflector (RSR), mirrors etc. for the laser delivery path, the process monitoring path, or the metrology beam path. The back-emitted radiation from the melt pool is switched between the process control module or the TEMPS module by a manually swapped, kinematically mounted RSR accessed by opening the top lid of the tower. Either a broadband reflector, or laser-line reflector is exchanged in the RSR position to switch operation between the process control module, or the TEMPS module. The optics tower is kinematically mounted to the optics bridge, which is rigidly attached to the granite base, however it is mechanically isolated from the vacuum chamber through a sealed, vacuum gasket flange.The process control module sits atop the chamber adjacent to the optics tower, and is the heart of the AMMT process monitoring for additive manufacturing research.

This module is essentially an enclosed optical breadboard which contains the laser fiber, beam expander, negative lens, and other bending mirrors or windows for the laser injection beam path to the optics tower. When the broadband reflector (700 nm to 1100 nm) is installed in the RSR position within the optics tower, this module also receives the back-emitted radiation from the melt pool which is further filtered to the 700 nm to 900 nm range with another beamsplitter. This 700 nm to 900 nm bandwidth can then be further split to various optical sensors, including high speed imagers or photodetectors.

This 700 nm to 900 nm bandwidth can then be further split to various optical sensors, including high speed imagers or photodetectors. The process monitoring module provides extra space for these different optical configurations and instruments to use the back-emitted light for researching various co-axial melt pool monitoring methods.The TEMPS module (not shown in Figure 1) contains the toroidal mirror, folding mirrors, mirrored field stop, and other components to support spectral emissivity and reflectance measurements. The mirrored field stop allows back emitted radiation to pass through an aperture to various instruments, while reflecting the surrounding image to a high speed thermal imager. This allows the imager to ‘see’ the thermal field around the aperture, and help discern the relative location on or near the melt pool that is passed through the aperture to the sensors.

This allows the imager to ‘see’ the thermal field around the aperture, and help discern the relative location on or near the melt pool that is passed through the aperture to the sensors. TEMPS instruments include Fourier transform infrared (FTIR) spectrometer, visible array spectrograph, and filtered radiometers. Although not in the same physical location as the TEMPS module located behind the chamber but not shown in Figure 1, an in-situ hemispherical reflectometer (also not shown in Figure 1) used only for TEMPS operation will be incorporated within the build chamber, and be automatically placed on or removed from the build plane.Structural/Metrology LoopUltimate positional and mechanical tolerance requirements of commercial LPBF systems may be dictated by the size of metal powder used (10 s of micrometers).

However, as a metrological instrument, AMMT structural design aimed to not only reduce potential sources of geometric error in AM builds, but to ensure stability and repeatability of measurements. Rules and concepts from precision machine design are apt for incorporation into AM machine design. Analysis of a manufacturing system with respect to the structural/metrology loop(s) enables a systematic approach to error identification and tabulation. A guiding principle in precision machine design dictates that a smaller, less complex, and closed structural/metrology loop provides improved precision [6].The structural loop, shown in Figure 2, consists of the mechanical coupling of rigid structures from the build layer to the scanner mirrors and laser beam path back to the build layer, with the granite base forming the central reference structure.

As mentioned in the previous report1149[1], the build bed, powder feed bed, and calibration sources are suspended in a gantry-type assembly from two rails on the granite base, and can be positioned along the X-direction with respect to the granite base using a ballscrew mechanism (not shown on the back side of the main carriage in Figure 1). This allows either the build bed or the calibration sources located within the carriage to be positioned directly under the optics tower, and accessed via the laser or co-axially aligned instruments. Using the CAD assembly, size and location of the carriage components were arranged to ensure the calculated center of mass is close to the mechanical center of the rails as possible.

Using the CAD assembly, size and location of the carriage components were arranged to ensure the calculated center of mass is close to the mechanical center of the rails as possible. As previously mentioned, the optics tower is mechanically decoupled from the vacuum chamber, but rests upon kinematic mounts on the optics tower bridge, which is then coupled to the granite base.YZGraniteGraniteCarriage Gantry + Z-stage BaseBuild Bed Arm/Stack AssemblydZdXrx, ryOptics Tower BridgeOptics Tower and Mounted GalvoMirrorsBuild PlaneBallscrewMountsFigure 2: Mechanical linkage and metrology loop. Left: Solid model assembly view (multiple components hidden for clarity).

Left: Solid model assembly view (multiple components hidden for clarity). Right: Schematic showing rigid structures, relative motion between rigid structures, and the metrology loop.Though concepts from precision machine design were incorporated into the structural design of the AMMT system, many concepts were not directly transferable, and there is room for further research into the combination of machine tool metrology and optical system metrology for AM systems, especially when process monitoring instrumentation is considered. The AMMT, apart from the research objectives outlined above, will also act as a platform for researching the effects of machine components on part quality, and identify methods and requirements for error compensation and error budgeting.1150Build Bed and Recoater AlignmentThe designs of two subsystems on the carriage are detailed here, which differ from common commercial LPBF system configurations.

Quality LPBF builds require consistent, smooth, and level powder layers, and thus require accurate alignment between the recoating plane and the build surface. The edge of the recoater blade forms a line in free space, and defines a plane when the recoater arm moves in the X-direction, shown in Figure 3 left. For uniform powder layer thickness, this recoater blade plane must be as parallel as possible to the build plate plane. Build layer thickness is then defined by the dz motion of the build arm motion stage. To parallelize two planes in free space requires two rotational adjustments with respect to one another. Many LPBF machines parallelize the build/recoater planes by implementing two rotational adjustments on the build plate, whereas we choose to separate the adjustment. Figure 3 left, demonstrates the relative motion and angular adjustments of the build plate and recoater blade. Not shown are the plate heating element, ceramic standoffs, or spring/felt seal assembly.

Not shown are the plate heating element, ceramic standoffs, or spring/felt seal assembly. The recoater blade is adjusted by a differential screw on one end, which rotates the blade about pivot bearings (angular flexures) on the other end. The recoater arm is positioned in δx by a motorized linear stage attached to the vertical wall on the granite, and held by an outrigger ball rail on the opposite side, also attached to the horizontal face of the granite.ZryRecoaterarm motion stageFulcrum for 3:1 leverrxdZdxBuild plateRecoaterblade angular adjustmentBuild plate angular adjustmentBuild arm motion stage Side-view of build plate servo-controlled levelling mechanismFulcrum for Build PlatexyzFigure 3: Left: Build plane and recoater plane parallelization adjustment coordinates. Note that some stiffening brackets and the outrigger ball rail connection are not shown. Right: Cutaway view of build arm showing internal linkage assembly for the automated build plate adjustment.

Note that some stiffening brackets and the outrigger ball rail connection are not shown. Right: Cutaway view of build arm showing internal linkage assembly for the automated build plate adjustment. An actuated pushrod rotates a 3:1 lever arm, which rotates the build plate’s free edge about a pivot bearing fulcrum.1151Provided the tight working space, an automated levelling mechanism was incorporated to align the build plate, shown in Figure 3 right. Again, certain elements such as the plate heater, standoffs, and spring/felt seal assembly are not shown. The five bar linkage mechanism transfers a linear displacement from a stepper motor with worm and rack gearbox, through a 3:1 lever arm (for mechanical advantage and improved positioning precision), to a pivoting build plate which rotates about the y-axis (ry).

Based on an estimated maximum distributed load, calculated design considerations included overall build plate stiffness, stiffness and strength of the pivot bearings, and detent torque of the motor.Optical System DesignThe driving design criteria for the process monitoring path is to create a 1:1 imaging system in the near infrared (NIR) which co-axially aligns an imager with the laser injection path, thus creating a stationary image of the melt pool. The 1:1 magnification was chosen so that a nominally 100 μm to 200 μm wide melt pool can be resolved between 10 pixels to 50 pixels, where typical Silicon (Si) based detector pixels are commonly between 4 μm to 10 μm. The near infrared spectrum was chosen for two reasons. Silicon based detectors, which are sensitive to visible and NIR, are less expensive and have a wider range of commercially available options.

The near infrared spectrum was chosen for two reasons. Silicon based detectors, which are sensitive to visible and NIR, are less expensive and have a wider range of commercially available options. Another reason NIR was chosen, as opposed to visible spectrum, is so lower melt pool temperatures near the expected nickel or steel alloy melt pool solidus temperatures (around 1500 °C) can be better detected, since the maximum greybody radiant emission at this temperature is approximately 1.6 μm. However, choosing longer wavelengths is limited by the Si detector sensitivity cutoff at approximately 1200 nm, and the need to avoid the 1070 nm laser wavelength, which would potentially damage the detector. For the optical design, an imager sensitivity at 850 nm was chosen.Multiple other sensor options and capabilities can be made possible within the process monitoring module.

For the optical design, an imager sensitivity at 850 nm was chosen.Multiple other sensor options and capabilities can be made possible within the process monitoring module. However, it was known early on that design requirements for the imager were most stringent and higher priority, and other sensors may be incorporated using various turning mirrors, beam splitters, and or hot/cold mirrors. Provided an imaging system with adequate spatial and spectral resolution is available, incorporation of other single-point detectors such as filtered photodetectors or pyrometers is trivial.Optical Modeling and Optimization of the Process Monitoring PathThe optical system consists of laser injection and process monitoring paths co-located and sharing some components, shown in the Figure 4 schematic. The laser delivery fiber is terminated with a 60 mm focal length collimator producing collimated light at approximately 1070 nm.

The laser delivery fiber is terminated with a 60 mm focal length collimator producing collimated light at approximately 1070 nm. This light is reflected by the wavelength separator, whose coating is designed to reflect 1070 nm light in a narrow band, and transmit light in the NIR. The reflected laser light propagates through a Linear Translating lens (LTZ), through the window and Converging Lens Pair (CLP), and is then reflected by the mirrors in the scanner and converges to a focused spot onto the build plane. In the reverse direction, NIR light emitted from the build plane reflects off the mirrors in the scanner, through the CLP and window, through the LTZ, through the laser reflector, and is then directed through Custom Imaging Lenses (CIL) that forms an image for a camera to view.

Fold mirrors and/or a custom relay lens can be used to direct the beam to be imaged to a more convenient location as needed.1152The laser injection path is required to generate a focused spot having a full-width half-max (FWHM) of approximately 100 μm at an arbitrary location within a 100 mm x 100 mm square area below the scanner. It is the function of the LTZ to maintain the focus by compensating for the optical path length change resulting from the scanner mirrors’ direction of the laser beam to the desired build plane locations. Commercial-off-the-shelf (COTS) parts for the LTZ and CLP are very good at maintaining this focus, but they were designed to work only over a very narrow wavelength band about 1070 nm.

COTS parts are inadequate because of two requirements: 1) it is desired to monitor and image light from a 12 mm region at the build plane emitted in the NIR with the best possible image quality (the “diffraction limit”) and 2) a focused, high power laser spot needs to be generated on the build plate at 1070 nm wavelength.Custom LTZCustom Imaging Lenses (CIL)ImageFold mirrorLaser reflectorLaserRemovable Splitter/Reflector (RSR)12 mm circular source object on build planeCustom Converging Lens Pair (CLP)ImageScanner and Fold Mirrors(Note: only one of three mirrors shown)Window(fused silica)Custom relay lenses(may or may not be used)To vacuum window and TEMPS toroidal mirror (when RSR is replaced with laser-line reflector)Figure 4: Schematic of ray tracing model used to optimize components for 1:1 imaging in process monitoring beam path.

Current beam paths shown are for AMMT operation, where the laser-line reflector is in the RSR position.To achieve diffraction-limited imaging performance in the NIR with 1:1 magnification in one direction as well as a focused laser beam at 1070 nm in the opposite direction, an “achromatic” solution is necessary, and custom LTZ, CLP, and CIL lenses are required. The lens surface curvatures and glass types were optimized using the ray-tracing optical design software based on a set of design criteria. The required operating wavelength bands, field-of-view, component temperature, and magnification all are included as constraints in the optimization routine.

The required operating wavelength bands, field-of-view, component temperature, and magnification all are included as constraints in the optimization routine. During the optimization process, the software automatically adjusts lens curvatures, glass types, and separation distances (the variables), and all within specified boundary constraints, resulting in simulated imaging performance characteristics (the objective functions).1153The result of the optimization process yields a solution for the imaging having the performance shown in Figure 5. On the left-hand side, a laser beam directed to a point located on the build plane at (30,30) mm is shown. In the middle of Figure 5, spot diagrams for light at 850 +/- 25 nm emitted by a point at (30,30) and (30, 24) mm are shown. A spot diagram is an analog of the geometrical Point Spread Function (PSF) and diffraction effects are ignored. It illustrates the geometric image blur for a point source.

A spot diagram is an analog of the geometrical Point Spread Function (PSF) and diffraction effects are ignored. It illustrates the geometric image blur for a point source. Superimposed on each spot diagram is the “Airy disk”, a contour whose radius represents the limit to the resolution when effects of diffraction are taken into account. Imaging can be no better than the diffraction limit. The radius of the Airy disk is a function of wavelength and system f-number (F/#) (Airy radius = 1.22⋅λ⋅F/#) and is shown to be about 16 μm in Figure 5. The geometric spot diagram shows that the extent of the blur for a point source lays within the Airy disk and thus the performance of the imaging is diffraction- limited. Another way to express the performance of the imaging is in terms of the Modulation Transfer Function (MTF), shown at the right-hand side in Figure 5. MTF is a direct measure of how well details in the object are reproduced in the image.

MTF is a direct measure of how well details in the object are reproduced in the image. Since the imaging is 1:1, linear coordinates on the build plane are unscaled at the image plane and the MTF plot directly represents the spatial resolution at the build plane. MTF is widely accepted as the most important criterion for judging image quality and is typically specified in terms of spatial frequency as “line pairs per mm” or “cycles per mm”. Diffraction effects limit the value of the MTF as the spatial frequency increases. As can be seen, the performance of the custom system results in an MTF that overlays the diffraction-limited MTF.40 mmModeled Build Plane(x,y) = (30,24) mmRelative AttenuationData for l= 825 nm to 875 nm(x,y) = (30,30) mmAiry disk contour is for 850 nmRadius = 1.22λ(f/#)Build point shown is (x,y) = (30, 30) mmXYFigure 5: Optical modeling and optimization results for 1:1 imager resolution.

Left: Schematic of physical model showing laser spot position coordinates. Center: Example spot diagrams showing point spread function (PSF) is encircled by Airy disk (thus diffraction limited performance) at center and 6 mm from center. Right: Modulation transfer function (MTF) showing spatial resolution.Build Process and Environment ControlThe AMMT control system design is process oriented and layer structured. The typical processes involved in AMMT can be divided into five major categories as shown in block diagram in Figure 6, and described as follows:1. Laser control – laser power and gating, scan path (galvo position) and LTZ control. 2. High speed monitoring – in situ process monitoring, experiment data acquisition, and real time feedback.3. Powder bed – build plate, powder feed, powder spread, and carriage motion and positioning.11544. Environment control – Build chamber preparation through vacuum purge, backfill, and inert gas circulation.5.

Environment control – Build chamber preparation through vacuum purge, backfill, and inert gas circulation.5. Environment monitoring – oxygen level, humidity, temperature, pressure, etc. monitoring.Processes 1-3 are related to the building and build monitoring processes and realized using National Instrument (NI PXIe) and Aerotech systems, with main operations programmed in LabVIEW. Processes 4-5 are environment related and implemented by an Allen Bradley Programmable Logic Controller (PLC), programmed in Studio 5000 by ladder logic.

Processes 4-5 are environment related and implemented by an Allen Bradley Programmable Logic Controller (PLC), programmed in Studio 5000 by ladder logic. The NI and PLC systems communicate to each other through an Open Platform Communications (OPC) server, where parameters in the PLC are published as global variables, read and written by Labview through the Datalogging and Supervisory Control (DSC) module.The PLC is constantly monitoring and will shut down the system, in particular the laser, and trigger an alarm in cases such as abnormal high temperature, high O2 level inside or low levels outside the chamber, chamber doors or panels open, etc.

Since the PLC runs independently from building process control (NI controller), this provides a secondary guard against potential safety hazards.Process CategoriesDevicesI/ORTOSHMIControl SystemLaser ControlNI SystemHigh-speedMonitoringPowder Bed EnvironmentControlPLC SystemEnvironment monitoringAerotechNPADLabviewVacuum PumpAllen Bradley CompactlogixController & ModulesStudio 5000CompressorPlatform & Recoater MotorsXY2-100OPCDSCNI FPGA & I/O ModulesAerotech RT/ A3200IR DiodeNI PXIe-1085 Labview RTValvesGalvo MotorsSensorsIPG LaserCameraFigure 6: Overview of the NIST AMMT control architecture.1155The block diagram is based on hardware, and much simplified and presented here to explain the architecture for AMMT control. Figure 6 can also be viewed as a layered structure from left to right.1. Device – Individual hardware to perform the tasks listed in process categories. 2. I/O – Input / Output.

Figure 6 can also be viewed as a layered structure from left to right.1. Device – Individual hardware to perform the tasks listed in process categories. 2. I/O – Input / Output. This is a heterogeneous system and in general this layer refers to the interface between system controller and individual hardware device or driver.3. RTOS – Real Time Operation System. NI PXIe-based system runs NI Linux Real-Time operating system, which also runs A3200 motion control Aerotech RTOS. On the RTOS, the PLC scans through all ladder logic in one clock cycle.4. HMI – Human Machine Interface, which can also be referred to as the application layer. A Windows-based personal computer (PC) is used, with Labview and Studio 5000 installed. The user interface (UI), developed in Labview, also resides on this PC. All data, including process feedback and environment monitoring, are eventually logged here.Environment Control ProcessA build process starts with build chamber preparation.

All data, including process feedback and environment monitoring, are eventually logged here.Environment Control ProcessA build process starts with build chamber preparation. The chamber is first vacuumed to the preset level (0.1 Pa or 10-3 torr), then back filled by inert gas (argon or nitrogen). The experiment can also be conducted directly in vacuum if required. Refill is controlled by the PLC proportional-integral-derivative (PID) loop formed by proportional valves and pressure sensors. After the desired pressure is reached, the oxygen level is checked and if it is higher than the desired value, the vacuum-backfill (inert gas) process is repeated until the O2 level is below the setpoint. The build cycle can then start.The build cycle (powder feed, spread, and laser melting) can be started with vacuum purge, or inert gas recirculation.

The build cycle can then start.The build cycle (powder feed, spread, and laser melting) can be started with vacuum purge, or inert gas recirculation. The circulation is powered by a 4 kW compressor, inert gas is pumped out of the chamber, filtered, cooled, and fed back into the chamber through three different ports into the chamber: 1) the air blanket, which is a custom-designed laminar flow nozzle directed right above the build plane, intended to uniformly remove vapor or melt pool ejecta, 2) the air knife, which is a high speed, thin flow layer located right below the scanner optics, intended to protect from contamination or deposition, and 3) the optics tower inlet, which maintains temperature of the mirrors and optics.After the build process finishes, the inert gas is pumped out, and the chamber is back filled with dry air.

The oxygen level warning sign will turn green only after the oxygen level reaches above 19.5%, to indicate it is safe to open the chamber door.Laser ControlFor LPBF AM, the laser beam is directed to the powder bed by a pair of mirrors driven by galvo motors. The galvos are limited rotational direct current (DC) motors with very fast response time, but are also very sensitive to noise. To eliminate transmission noise, an industrial standard digital communication protocol xy2-100 is used to transfer control signal (position) from the field programmable gate array (FPGA) to the galvo driver. The xy2-100 packages each position into a 20 bit ‘word’ and transmits it at 100 kHz per word (2 Mbps). Figure 7 shows the scheme for AMMT laser control. The scan path is generated on the controller PC together with laser power level according to the predefined scan strategies, and converted into xy2-100 format on the FPGA.

The scan path is generated on the controller PC together with laser power level according to the predefined scan strategies, and converted into xy2-100 format on the FPGA. The digital signal is transmitted as differential pairs and converted back to analog voltage by the1156digital to analog (D/A) receiver. It is then fed into the galvo driver as command signal, and form a closed-loop control with the position feedback from the galvo encoder.The standard xy2-100 protocol is only for positions (x, y, and z axis). The z-axis is typically relegated for controlling an LTZ lens. However, the AMMT utilizes this axis for laser power control to ensure high speed synchronization, as shown in Figure 7. The laser power level is also updated at 100 kHz, and is synchronized to each scanning step by introducing a constant, calibrated delay (laser response is much faster than galvo).

The laser power level is also updated at 100 kHz, and is synchronized to each scanning step by introducing a constant, calibrated delay (laser response is much faster than galvo). Since the LTZ position strictly depends on galvo x-y position values, and not on commands from the controller PC, it is controlled through a lookup table programmed on the FPGA, which is based on angular measurements determined from optical ray tracing. This lays down the foundation for laser path planning and power control strategies.Figure 7: Laser path and power control.Prototyping SystemCertain design considerations for the AMMT/TEMPS required physical testing and evaluation of components, in particular to determine attributes that cannot be modeled using ray tracing software. In addition, the anticipated complexity of the monitoring and control system warranted attention prior to final system construction.

In addition, the anticipated complexity of the monitoring and control system warranted attention prior to final system construction. In order to conduct these tests and initiate controller programming, the prototyping system was built which incorporates the laser, optics for the laser injection, galvo mirrors and driver system, and similar COTS components that will go into the process control module. Figure 8 shows a solid model schematic of the prototyping system with highlighted beam paths. Grantham et al. described several initial characterization tests, including laser spot size measurements, temporal laser control, and thermal load testing of various components to determine heating rates due to continual laser operation [3].

In the final AMMT/TEMPS facility, a prototyping system will exist along-side the full AMMT system with replicate components and similar capabilities to allow various testing and development before implementation on the AMMT.Multiple other tests are currently being performed on the prototyping system, and are published elsewhere in this conference proceedings. These include system performance and control tests to evaluate the scan speed, acceleration, laser on/off timing, and positional accuracy of the laser spot. Results from these tests determined that the COTS galvo controller and command software were inadequate for our purposes, which initiated development of a custom design.

Results from these tests determined that the COTS galvo controller and command software were inadequate for our purposes, which initiated development of a custom design. Other tests include basic process mapping tests of single weld tracks, and simple scan patterns to provide qualitative comparison to commercial LPBF systems.1157BeamsplitterNegative lensScannerFiber laser beam850 nm beam for imagingLeaked Fiber laser beam(power monitor)CollimatorFold mirrorLaser windowArpurge boxImaging lensFocusing lensSampleFigure 8: Schematic showing physical location of components within the prototyping system. The prototyping system is built on an optical table surrounded by a laser safety curtain.One of the more important tests was to determine the level of optical throughput initiating from an incandescing melt pool and passing to the co-axially aligned sensors.

Although one can calculate the expected optical throughput, adequate knowledge of the true melt pool temperature and emittance is very limited, therefore physical testing was necessary. Essentially, these tests were to determine at what integration time could the imager ‘see’ the melt pool. Though the custom optics for optimized resolution were not yet available, comparable signal from final optical design throughput is well approximated by the prototyping system.The imager tested had 5.5 μm/pixel pitch, quantum efficiency of 63% (measured at 545 nm), and used the 850 +/- 25 nm bandpass filter. While scanning on bare 17-4 stainless steel (no powder) at 800 m/s scan speed and 200 W laser power, the center of the melt pool just saturated the imager (reached 4096 digital levels at 12-bit A/D) at an integration time of 100 μs, shown in Figure 9. During the 100 μs, the melt pool traveled 80 μm, or the approximate equivalent of 15 pixels.

During the 100 μs, the melt pool traveled 80 μm, or the approximate equivalent of 15 pixels. However, given the fact that the melt pool is stationary within the image, any motion blur will stem from perturbations in the melt pool size or intensity rather than its motion. Although the true melt pool size (defined by the liquidus/solidus boundary), approximated temperature, or other absolute measures cannot yet be determined, these tests indicate that there is adequate signal level to move forward.1158200 mmCamera Signal (Digital Levels)Figure 9: Melt pool image from co-axial imager in the prototyping system. Initial imaging tests were used to determine imager signal level vs. integration time.ConclusionsThis report reviewed the higher-level design of the AMMT/TEMPS system currently being constructed and tested at NIST.

integration time.ConclusionsThis report reviewed the higher-level design of the AMMT/TEMPS system currently being constructed and tested at NIST. The system is essentially an LPBF additive manufacturing system, with a plethora of primarily optically-based sensor systems to enable high fidelity and well- characterized analysis of the melt pool and heat-affected zone. The two overarching purposes (additive manufacturing and materials optical property research) meet the combined interests of the NIST Engineering Laboratory (AMMT) and Physical Measurement Laboratory (TEMPS), respectively. The overall design applies concepts of modularization and accessibility to allow users to work to more easily on the instrument, and access and configure various modules separately from the main system.

In addition, overall design flexibility and reconfigurability allow new experiment designs to be more easily incorporated, or various optical beam paths to be accessed for material property characterization, or AM monitoring and control. Foremost, full controllability and characterization enable users to fully define the LPBF process parameters to a high precision, and know how the system responds.References[1] Vlasea M, Lane B, Lopez F, Mekhontsev S E and Donmez M A 2015 Development of powder bed fusion additive manufacturing test bed for enhanced real time process control Proceedings of the Solid Freeform Fabrication Symposium (Austin, TX) p 527[2] Dunsky C 2014 Process monitoring in laser additive manufacturing Ind. Laser Solut.

SIGN, FABRICATION AND EVALUATION OF NEGATIVE STIFFNESS ELEMENTSLia Kashdan, Dr. Carolyn Seepersad, Dr. Michael Haberman, Dr. Preston S.Wilson Mechanical Engineering Department University of Texas at Austin, Austin, TX, 78712AbstractRecent research has shown that constrained bistable structures can display negative stiffness behavior and provide extremal vibrational and acoustical absorptive capacity. These bistable structures are therefore compelling candidates for constructing new metamaterials for noise reduction, anechoic coatings, and backing materials for broadband imaging transducers. To date, demonstrations of these capabilities have been primarily theoretical, because the geometry of bistable elements is difficult to construct and refine with conventional manufacturing methods and materials.

The objective of this research is to exploit the geometric design freedoms provided by selective laser sintering (SLS) technology to design and construct constrained bistable structures with negative stiffness behavior. The static and dynamic behaviors of resulting bistable structures are experimentally investigated. Initial bistable designs and test results are presented in this paper.1. IntroductionNegative stiffness elements have been identified as unique mechanisms for enhancing acoustical and vibrational damping. Examples of negative stiffness mechanisms include mechanical systems with negative spring constants (Prasad, 1991) and materials with negative moduli (Lakes, 2001a, 2001b). Negative stiffness elements contribute to damping behavior because they tend to assist rather than resist deformation as a result of internally stored energy (Lakes, 2001a, 2001b).Figure 1: Stages of snap-through movement of a buckled beam. (i) The beam is in an initial stable buckled state.

(i) The beam is in an initial stable buckled state. (ii) A force is applied to the spring and sends the beam to the meta-stable state. (iii) The spring is constrained, and a slight change in displacement sends the beam to (iv) the second stable buckled state. Image from Haberman, 2007.A simple negative stiffness structure can be created from a buckled beam, as illustrated in Figure 1. Such systems are bistable when unconstrained and include one metastable equilibrium point. Figure 1 illustrates a buckled beam with pinned ends moving though the three points of equilibrium. In (i) the beam is in the first stable equilibrium state and when a force is applied, it is pushed to the metastable state (ii-iii) and through to the second stable state (iv).1D Example: Post-buckled BeamForce vs. Transverse DisplacementNormalized Transverse Force of a Beam with Axial Load vs.

transient displacement. Points (1), (2) and (3) match those in Figure 2 as well as the legend for the unbuckled, critically buckled and completely buckled states. Image from Haberman, 2007.The force versus transverse displacement of a buckled beam is plotted in Figure 2, and the strain energy versus transverse displacement is plotted in Figure 3. In both figures, the behavior of a completely buckled beam is plotted alongside the behavior of an unbuckled beam. As illustrated in Figure 3, the equilibrium positions of the buckled beam correspond to local minima and maxima of the strain energy curve. Since the stiffness of the beam corresponds to the spatial derivative of its strain energy, the buckled beam exhibits negative stiffness over the interval indicated in Figure 2.

Since the stiffness of the beam corresponds to the spatial derivative of its strain energy, the buckled beam exhibits negative stiffness over the interval indicated in Figure 2. Therefore, negative stiffness can be elicited by applying a transverse force to the center of a beam in state (1) or state (3), each of which is characterized by a strain energy minimum, and transiting the metastable state (2). The buckled beam is a simple 1D example of a bistable structure, and negative stiffness behavior is implicit to constrained bistable elements of any configuration (Prasad, 2006; Qiu, 2004).This behavior has been confirmed by experimental studies of buckled tubes and carbon nanotubes which found decreased force with increasing strain response and higher damping levels for buckled tubes relative to unbuckled tubes under sinusoidal forcing (Lakes, 2000; Yap, 2007).

These findings confirm that buckled beams are negative stiffness elements under specific conditions and that they can therefore be used for damping purposes. However, it is difficult iteratively design and tune negative stiffness elements with conventionally manufacturing methods. Our research is focused on designing, tuning, and testing negative stiffness elements by fabricating them with SLS technology. This research is part of a larger project to test these mechanisms at macroscopic scales and then miniaturize them to create metamaterials or waveguides with very high damping capacity.The process of using additive manufacturing technologies to create structures or materials that provide acoustic or vibrational damping has not been studied extensively.For example, SLS has been used to produce a passive destructive interference acoustic absorber (Godbold, 2007).

The absorber minimized sound at specific frequencies by sending acoustic waves through two channels of different lengths thereby creating destructive interference at the exit. However, destructive interference is very different from the negative stiffness mechanism investigated in this research, and negative stiffness has not been investigated with SLS technology. SLS offers the benefit of increased through design design freedom for fine-tuning permutations, and consolidating parts, relative to conventional manufacturing techniques.the geometry, rapidlyiterating2. Research MethodologyThe methodology for the research study is presented in Figure 4. This paper follows the methodology through the initial macroscale design phase and reports results of material characterization and design, fabrication, and testing of macroscale designs. Future work will involve miniaturizing our results into smaller-scale and multiscale designs and tuning those designs for specific examples.

Task Clarification and Overall Experimental DesignA simple apparatus was designed for testing and tuning the negative stiffness behavior of a buckled beam. As shown in Figure 5, the apparatus consists of a beam connecting two end plates. The endplates are connected together with bolts that can be tightened to pre-buckle the beam. By adjusting the relative tightness of the bolts, the beam can be initialized into an unbuckled state, a metastable state, or a fully buckled state. The spring serves to constrain the specimen; its stiffness is tuned to stabilize the beam in its metastable state when the bolts are tightened to the appropriate length. As described in Section 5, the apparatus is placed on top of a shaker table, with accelerometers attached to the beam and the base of the structure to measure the relative amplitude of the beam’s motion and the transmissibility of the beam.

The amplitudes of resonance peaks from the pre-buckled beam are expected to be much lower than those of a beam in its free state, demonstrating the dampening effect of the bistable element. The negative stiffness behavior of the apparatus is very sensitive to the material properties of the beam and the spring. Since the apparatus is being constructed with SLS technology and Nylon 11 material, for which various material properties are not widely documented, it is important to begin the study with material testing.Figure 5: SolidWorks model of a buckled beam system4. Material CharacterizationStatic materials testing is a common and relatively straight-forward method for obtaining material properties of sintered parts. However, it may not always be the most appropriate method when parts are used in dynamic applications.

However, it may not always be the most appropriate method when parts are used in dynamic applications. Because the stiffness properties of materials such as Nylon 11 change with strain rate, it is very important to employ dynamic tests to determine properties at a specific frequency. These tests are of added value to this research because limited material property data is available in the literature for sintered Nylon 11. For the purposes of this study, Young’s modulus, shear modulus and loss factor are determined using dynamic tests and static tests. All parts aremade using the 3D Systems Sinterstation HiQ SLS machine and the build parameters outlined in Appendix A.4.1.Shear ModulusFor this system, shear modulus must be determined to calculate spring coefficients for the spring element that will be included in the test model. The shear modulus is related to the spring coefficient using!"#$%&’ ()*+ ,&-.

The shear modulus is related to the spring coefficient using!"#$%&’ ()*+ ,&-. ,(1)where G is the shear modulus, d is the coil thickness, n is the number of active coils and D is the mean coil diameter. The shear modulus is determined using this relationship for both static and dynamics tests of sintered spring elements.For the static test, two sets of springs (one set seen in Figure 6) with varying coil thicknesses and outer diameters (see Table 1) are produced with one set sintered vertically and the other horizontally. The springs are each loaded in an Instron 3345 Tension Tester machine to obtain force versus displacement curves, thereby directly obtaining kspring by finding the slope of the resulting curves.Figure 6: Three types of spring designs used in static testingTable 1: Dimensions of springs for static and dynamic testsnote that n refers to the number of active coils, not total coils.

In the case of the springs used in this system, n is the total number of coils minus one.Static!Test!"!tested!at!strain!rate!of! 5!mm/s!Dynamic!Test!–!tested!from!5!to!60!Hz!! /! 0!n* 9 8.5d (mm) D (mm) 2.75 2.765.00 9.881 2n* 2.5 4.5d (mm) D (mm) Mtot (kg) # Springs 4.6 4.149.9 20.20.12 0.112 31!8.53.3210.18The parts fabricated for the dynamic test can be seen in Figure 7. These parts are designed to be simple mass-spring systems consisting of two or more springs to insurethe stability of the mass during forced vibration (see Table 1). This configuration is chosen for its simplicity and its well-known mathematical model.Figure 7: Two different mass-spring systems used for dynamic testing.

This configuration is chosen for its simplicity and its well-known mathematical model.Figure 7: Two different mass-spring systems used for dynamic testing. The system on the left contains three springs and the one on the right has two springs in a “double helix” configuration for stability of the massAccelerometersSignal!ConditionerMass!– Spring! SystemPower!AmplifierShaker! TableComputerVector!Signal! AnalyzerFigure 8: Schematic for dynamic shear modulus testThe test used to determine the dynamic shear modulus is a standard vibration transmissibility test schematized in Figure 8. The spring-mass element is fixed to the top of a shaker table, which is driven by a frequency-swept sinusoidal voltage signal. During the sweep, the acceleration of the base of the sintered part and the top of the mass is measured using two matched accelerometers (PCB 333B30). A ratio of the mass acceleration to the base acceleration yields the transfer function of the system.

A ratio of the mass acceleration to the base acceleration yields the transfer function of the system. The peak of this function occurs at the resonant frequency of the system, which for the case of light damping is related to the spring constant by2& ( 3!"#$%&’5 , 4where !n is the natural angular frequency at resonance, and M is the mass of the blocks shown in Figure 7 plus the accelerometer and 1/3 the total mass of the springs. The measurements of the shear modulus from both these tests are given in Table 2.(2)Table 2: Results from shear modulus static and dynamic tests. Six springs were each tested ten times for the static tests, while two mass-spring systems were each tested five times.Average Standard DeviationG Measurements Static (Pa) 4.77x108 1.38 x108Dynamic (Pa) 6.44 x108 8.04 x107Results for the static test are very close to the published values of shear modulus for molded or extruded Nylon 11: 4.38 x 108 – 4.63 x 108 Pa (Granta, 2008).

The dynamic test results are outside this range. This result is likely due to the fact that Nylon 11, being a polymer, has strain rate dependent (i.e. frequency dependent) moduli and that the test measured the modulus at the resonant frequency.4.2.Young’s ModulusThe Young’s modulus of sintered Nylon 11 is also found using both static and dynamic methods. The static method used a tensile testing apparatus and follows ASTM D638, including strain rates specified therein. As Young’s modulus is highly rate dependent, simply determining this value from tensile testing may not provide adequate information, especially for systems that operate over a range of frequencies. It is for this reason that the material is also tested under dynamic conditions, specifically using the log decrement method for a cantilevered beam (Jones, 2001).Three sets of rectangular beams are designed with a range of dimensions and built in the 3D Systems Sinterstation HiQ SLS.

The test beams are clamped to a table (as seen in Figure 9) with an accelerometer affixed to the end of the beam.Acceler ometerNylon!11!BarTableClampSignal!ConditionerComputerOscilloscopeFigure 9: Schematic for Young’s modulus dynamic testThe beam is struck on the end to induce motion primarily at its resonant frequency. Output from the accelerometer, an example of which is shown in Figure 10, is captured on the oscilloscope and transferred to the lab computer for analysis. The resonantfrequency of the beam is a function of its length, thickness, density, and Young’s modulus. Multiple tests can therefore be run on each beam by shortening the free length in order to obtain modulus information for a range of frequencies. The resonant frequency of this system can be determined using a discrete Fourier transform of the time signal and the Young’s modulus is then calculated as follows (Thompson, 1987):6 (9 :;<;=.

The resonant frequency of this system can be determined using a discrete Fourier transform of the time signal and the Young’s modulus is then calculated as follows (Thompson, 1987):6 (9 :;<;=. >?78.Note that Mtot includes the entire mass of the accelerometer plus 0.23 times the mass of the beam, I is the area moment of inertia of the beam, and L is its length.Figure 10: Sample output from Young’s modulus dynamic testOver two hundred trials are run with this setup, and a chart of the resulting Young’s modulus calculated from the tests is presented below in Figure 11. The bin numbers underneath each column represent the highest value within that column.

No correlation to frequency could be determined for these beams as the frequencies sampled are mostly too low (less than 100 Hz), however the Young’s modulus would be expected to increase significantly at higher frequencies.(3)25.000%20.000%a t a D ! f o %!15.000%10.000%5.000%0.000%Yooung's!Moduluus!(MPa)Figure 11: RResults from thee Young’s moduulus test in histogram formT The expected 310 MPa (G to 13 ulus with ten modu just below falls rials often d mater for six sinte data dard deviatio stand he values fo and th and the fact t size a The advantag T iple tests to multi ulus at the n modu le tests that tensil sonic tests ar ultras sonic testing ultras oyed at low empl MHz) (Jones, 20 Md range of the Granta, 2008 n samples is the expecte display some ered Nylon 1 on of 172 M ound using th that the dyna ge of the dy o be perform natural frequ are limited t re another co g is not the w frequenciese Young’s m ).

The avera s 1220 MPa ed range. A e reduction i 11 bars resu MPa, a signif he log decre amic sample ynamic test med with a uency of the to the strain ommon meth best choice s as they stifmodulus for m age for the a, with a stan A lower mod in stiffness d ulted in an av ficantly lowe ement metho es are tested is its nond single bar. bar, which n rates set in hod for deter for the cha ffen significxtruded Nylo molded or ex ally determi experimenta tion of 97.7 ndard deviat is expected dulus value sity. Note th due to poros ulus of 1060 verage modu n both mold er value than due to the sm od possibly d e. at resonance nature, whic destructive n also determi The test a de more use could provi standard. It n the ASTM ung’s modul rmining You n of plastics aracterization rasonic frequ cantly at ultron 11 is 124 ined Young’ 7 MPa, whic d as sintere hat tensile te 0 MPa, and ded Nylon 1 maller sampl40 ’s ch ed st a 1 lech allows fo ines Young’ eful data tha t is noted tha lus.

Howeve s that will b uencies (0.5or ’s an at er, be -2001).4.3.Looss FactorKnowledge o K mines the da determ ame log dec the sa factor is fo loss ying transien decay at curve is re of tha termine the l to detof the loss fa amping of th crement data und by fitti nt response o elated to the loss factor aactor, ", of s he material. a used to de ing a logari of the beam logarithmic as follows (Josintered Nylo Convenient etermine the ithmic curve m, such as the c decrement o ones, 2001):on 11 is vita tly, the loss e Young’s m e to the pea e one shown of the responal to this stu factor can b modulus. Spe aks of the e n in Figure 1 nse, #, whichit udy because be found from m he ecifically, th ly exponentiall pe 10.

1100rom the mat s system. Fo ne side and is axially lo pring to the his configur due to the n ign permuta atio for the b in the beam and vibrated n of the syste fixed ends o length of 15terial charact or the system d the beam c oaded, the sp otherwise m ration will egative stiffn ation used p beam cross s m to its met , which prod em, as show on the beam. 55 mm andterization ar m under test, center point pring impart metastable st display en fness of the b pinned joints ection dimen tastable stat duced signif wn in Figure The new de a spring wire used to de a constraini on the othe ts the force o ate when th nhanced dam beam. s for the bea nsions, and h te. During e ficant noise 13, to inclu esign has be ith spring coesign and te ing spring is er, as shown on the beam he beam is b mping compte st a complet attached to a n in Figure 5 5. by m and thereb is buckled. It i he pared to tham, did not had a spring excitation, th in the data.

by m and thereb is buckled. It i he pared to tham, did not had a spring excitation, th in the data. ude a stiffer eam thicknes onstant of ahave a larg g that was to he pin joint These issue spring, large ss of 2.5 mm approximatelge oo ts es er m, ly(4)Figure 13: The buckled beam system made of Nylon 11The left wall of the system is free to slide towards the right to compress the structure. Threaded rods (not shown) were inserted through the left and right end plates of the part and were used to progressively compress the system by small increments. Holes in the base allow for direct attachment of the part to the shaker table (Labworks Inc. model ET- 139).

Holes in the base allow for direct attachment of the part to the shaker table (Labworks Inc. model ET- 139). The vibration transmissibility of the apparatus is then determined with the experimental setup shown in Figure 14, which is nearly identical to the test used to determine the shear modulus.AccelerometersBistable!SystemSignal!ConditionerShaker! TablePower!AmplifierComputerVector!Signal! AnalyzerFigure 14: Schematic of dynamic buckled beam testVibration transmissibility tests are run at multiple levels of axial pre-strain. The tests begin with the beam in a completely uncompressed position and the beam is progressively pre-strained until the beam is in a stable buckled equilibrium position, after passing though the metastable position. Each round of testing begins by compressing the beam by approximately 0.1 mm. The shaker table then excites the system in a sweep from 30 to 200 Hz, a range that contains the first resonance peak.

Each round of testing begins by compressing the beam by approximately 0.1 mm. The shaker table then excites the system in a sweep from 30 to 200 Hz, a range that contains the first resonance peak. Several representative transmissibility curves are shown in Figure 15. Table 3 matches each curve in Figure 15 with an exaggerated and approximate image of the pre-stressed beam before the shaker test began.Increasin ng g buckling765342 21Figure 15: Resp Fponse of systemduring increassed buckling. TThe curves are nnumbered for fuuture referencee.The results cl T ns higher val attain itude of the ampli ence of the prese dening of th broad increased ne with een more cle be se malized frequ normlearly show lues of nega e transfer fu buckled b he resonance egative stiffn early in Figu uency, therebthat as the b ative stiffnes unction at re beam increa e peaks is a ness. The red ure 16, whic by aligning thbeam becom s, the resona esonance.

The red ure 16, whic by aligning thbeam becom s, the resona esonance. T ases the ove another indic duced amplit ch displays he resonantmes more com ance frequen This second erall dampi cation of in tude and bro the transmi peaks.mpressed an ncy decrease d point indic ing of the ncreased syst oadening of t ssibility asm nd the system he es as does th he cates that th he system.

Th ng tem dampin an the peaks ca of a function oNormalized!Trannsmissibility!oof!Negative!Stiffness!System4030) B d ( ! !20| X | / | Y |100"10"204 0.40.61.2 1 1 quency!(f/fn) ormalized!Freq0.8No1.4Figure 16: FFrequency normmalized responsee of system durring bucklingP eak 1, the u dB w while peak 7unbuckled re 7, the most besponse obta buckled respains a maxim ponse obtainmum amplit ns a maximtude of appr mum amplituroximately 3 ude of 19 dB37 B.This is equivalent to an 87% decrease in damping which is notable considering that the total compression of the system is 0.772 mm. Furthermore, these negative stiffness systems can also be used to tune the natural resonance of a system by increasing or decreasing the amount of buckling accordingly.

Furthermore, these negative stiffness systems can also be used to tune the natural resonance of a system by increasing or decreasing the amount of buckling accordingly. This experiment demonstrates that negative stiffness systems are remarkably effective as damping agents.Table 3: This table outlines the actual compression during each stage in Figures 15 and 16. The image shows the exaggerated stages of deflection of the beam at each point. At peak 6, the beam reaches the end of its metastable condition and becomes completely buckled.Peak NumberTotal Amount of Compression (mm)Exaggerated Beam Deflection1020.04530.21340.28550.40960.60470.7726. ClosureThe goal for this project is to ultimately develop metamaterials that can achieve higher damping than currently available materials, via negative stiffness behavior.

ClosureThe goal for this project is to ultimately develop metamaterials that can achieve higher damping than currently available materials, via negative stiffness behavior. Towards this goal, macroscale proof-of-concept tests were conducted on a bistable element, specifically a pre-buckled beam, that has been shown theoretically to exhibit negative stiffness behavior. The apparatus was built with SLS technology and Nylon 11 material. A preliminary design for the apparatus was created, and material properties for sintered Nylon 11 were measured using both static and dynamic tests. Using those properties, the initial design could be tuned to obtain the correct damping behavior and fabricated with SLS. The results from dynamic experiments indicate that a simple negative stiffness system has the ability to increase damping, as resonance peaks for the pre-buckled beams (i.e., bistable elements) are lower in amplitude and broader than those for unbuckled beams.

The next step will be implementing more complicated negativestiffness systems, in two and three dimensions, and with composite materials. From there, the ability to use negative stiffness inclusions within materials will be explored.AcknowledgementThe authors would like to acknowledge an internal research and development grant from Applied Research Laboratories at the University of Texas at Austin for funding this research.ReferencesGodbold, O. B., R. C. Soar, et al. (2007). "Implications of Solid Freeform Fabrication on Acoustic Absorbers." Rapid Prototyping Journal 13(5): 298-303.Granta (2008), CES EduPack.Haberman, M. (2007) “Design of High Loss Viscoelastic Composites through Micromechanical Modeling and Decision-Based Materials Design,” PhD Thesis, G.W.

sign, Fabrication and Testing of Graphite Bipolar Plates for Direct Methanol Fuel CellsKaushik Alayavalli and David L. Bourell,Laboratory for Freeform Fabrication, The University of Texas at Austin,Austin, Texas 78712AbstractDirect Methanol Fuel Cells (DMFCs) are clean energy devices that convert chemical energy into electrical energy. The bipolar plate component of the DMFC is required to be fluid impermeable to prevent fuel leakage and electrically conductive to collect the electrons produced within the cell. Graphite possesses the properties of high electrical conductivity, low weight and resistance to corrosion that make it an attractive material for bipolar plates. However, the poor mechanical properties of graphite lead to prohibitive machining costs.

However, the poor mechanical properties of graphite lead to prohibitive machining costs. The indirect Selective Laser Sintering (SLS) process, involving the laser sintering of graphite powders mixed with a phenolic resin binder, offers the advantage of complex part production and testing of prototype bipolar plates in short times. The current research deals with the production of bipolar plates for DMFC's by utilizing indirect SLS followed by post processing steps.1. IntroductionDirect Methanol Fuel Cell (DMFC) technology is an attractive option as an alternative power source due to the theoretical energy density of methanol which is higher than current Li ion battery technologies and also hydrogen fuel cells. Moreover, methanol fuel is easy to handle and transport, making it a safe and viable means for portable energy solutions.

Moreover, methanol fuel is easy to handle and transport, making it a safe and viable means for portable energy solutions. However, issues such as methanol crossover [1] from the anode to the cathode side of the fuel cell and unstable electrocatalysts [2] hamper this technology from achieving widespread use.A DMFC functions as a power source by converting chemical energy into electrical energy. This is accomplished by supplying methanol directly to the anode of the fuel cell where it is oxidized to CO2 (Figure 1). The H+ ions are transported across a polymer electrolyte membrane to the cathode where they react with oxygen to form water.

The H+ ions are transported across a polymer electrolyte membrane to the cathode where they react with oxygen to form water. A typical DMFC stack is comprised of repeating units of a membrane electrode assembly (catalyst coated electrodes on either side of a polymer electrolyte membrane), diffusion layers, sandwiched between bipolar plates that supply and distribute the fuel, methanol at the anode and oxygen at the cathode.Bipolar plates serve the purposes of fuel distribution, current collection, separation of the membrane electrode assemblies and bearing of the clamping force holding the fuel cell stack together. They are the most significant components of a fuel cell stack in terms of weight and cost of production [3]. Graphite has been the standard material of choice for bipolar plates due toits high electrical conductivity and low mass.

Graphite has been the standard material of choice for bipolar plates due toits high electrical conductivity and low mass. However, machining fine features such as flow channels and fuel manifolds on graphite plates is expensive since graphite is brittle and has poor mechanical properties [4].Figure 1. Schematic representation of DMFC working principle.Selective laser sintering (SLS) is a rapid prototyping process that offers the benefit of producing plates with features of significant complexity in a single step. Indirect SLS of graphite bipolar plates involves mixing graphite powder with finely milled phenol formaldehyde (“phenolic”) powder and using a laser to scan individual layers of a .stl file into a solid 3- dimensional object.

Due to the nature of the indirect SLS process, which involves the melting of the phenolic binder by the laser to bind the graphite powder together, the ‘green’ SLS parts produced are porous and require post processing for the improvement of electrical properties and complete densification. The post processing steps include heating the green parts in a vacuum furnace to burn off the phenolic binder and infiltrating the porous ‘brown’ part with a suitable infiltrant to densify and strengthen the part. Bipolar plates made by indirect SLS offer the advantages of complex part production in a single manufacturing step. The functions of bipolar plates include distributing fuel and conducting electrons from the stack into the external circuit.

The functions of bipolar plates include distributing fuel and conducting electrons from the stack into the external circuit. However, due to the nature of the indirect SLS process, the as-produced (green part) plates are porous and possess low electrical conductivities.The performance of the DMFC is dependent on optimal fuel distribution across the catalyst layer as well as the presence of detrimental processes such as methanol crossover which occurs when methanol from the anode side, permeates through the membrane to the cathode side where it poisons the catalyst sites, reducing the overall efficiency of the cell. Optimization of the flow field design in the bipolar plate has been shown to affect the efficiency of the fuel cell. The current research work deals with the issues related with the production of highly conductive, fluid impermeable graphite parts by indirect SLS and the optimization of the flow field design to improve the operating efficiency of the DMFC.2.

Bipolar plate production by Indirect Selective Laser SinteringThe process of producing bipolar plates by indirect SLS as shown in Figure 2, involves the production of a green part by the indirect SLS of a mixture of graphite (~ 70 microns, GS-150E, Graftech Inc) and phenolic (~ 11 microns, GP-5546, Georgia Pacific) powders in a 70 wt% to 30 wt% ratio, in a Sinterstation 2000 [5]. The porous green part is then heated in a vacuum furnace at temperatures greater than 1200 ºC to carbonize the phenolic binder, thus producing a ‘brown’ part. Phenolic is a “high ash” polymer, producing significant residual carbon after thermal dissociation. This carbon is effectively a high-temperature binder of the graphite powder. The parameters used for the SLS process are: fill laser power (10 W – 20 W), outline laser power (4 W), powder layer thickness (0.1016 mm), laser scan spacing (0.0762 mm), laser scan speed (1.524 m/s).

Based on Differential Scanning Calorimetry (DSC) results of the phenolic powder, the part bed and feed bins were maintained at 65 ºC and 47 ºC, respectively, to ensure that the phenolic powder bed that was not laser sintered remained stable without being thermally affected. Post processing of this brown part includes infiltration with low viscosity cyanoacrylate glue [6] that seals the pores, thereby making the part fluid impermeable, and polishing of the surfaces to expose the underlying electrically conductive graphite matrix.Powder Preparation & MixingIndirect SLS of Green PartVacuum Furnace CarbonizationLow viscosity polymer infiltrationRoom temperature cure/ PolishingMaterial Properties Characterization and TestingFinished Bipolar PlateFigure 2. Flow chart of the experimental process of bipolar plate production by indirect SLS3.

Flow chart of the experimental process of bipolar plate production by indirect SLS3. Results and Discussion3.1 Effect of Carbonizing Temperature on Electrical ConductivityThough graphite has high electrical conductivity (~1000 S/cm), the phenolic powder used as a binder is electrically insulating, thereby rendering the green part produced by indirect SLS, a poor electrical conductor. Studies of phenol formaldehyde resin have shown that the electrical resistivity is highly dependent on the pyrolyzation temperature and that there is a change on the order of about 102 between treating at 800 ºC and 1200 ºC [7]. XRD was performed on the neat graphite and phenolic resin before and after pyrolysis at temperature ranging from 600 ºC to 1200 ºC. The results show that the carbon product of the phenolic burnout at 600 ºC is amorphous in nature and therefore may be contributing significantly to the drop in electrical conductivity of the part.

The results show that the carbon product of the phenolic burnout at 600 ºC is amorphous in nature and therefore may be contributing significantly to the drop in electrical conductivity of the part. As the carbonizing temperature increases, a noticeable sharpening of thepeak at ~ 43º 2θ is observed (Figure 3), which implies an increase in graphitic order of the carbon product which in turn signifies an increase in electrical conductivity.4 3Figure 3. X-ray diffraction patterns of phenolic burnout and neat graphite powder at varying carbonization temperatures.A four probe setup was used to measure the electrical conductivities of the heat treated brown parts and the results plotted as a function of the pyrolyzation temperature. The resulting graph (Figure 4a) corresponds to literature (Figure 4b) on the pyrolysis of phenolic resin [7].

The resulting graph (Figure 4a) corresponds to literature (Figure 4b) on the pyrolysis of phenolic resin [7]. This increase may be due in part to the conversion of amorphous carbon into a semi-crystalline phase and also a reduction in the porosity of the phenolic carbonization residue [6].Figure 4. (a) Conductivity vs. Pyrolysis Temperature(b) Resistivity vs. Pyrolysis TemperatureBhatia et al. J. Mater. Sci, 19, 1022 (1984)3.2 Bipolar Plate TestingFor purposes of testing and comparison, 3 sets of bipolar plates (anode and cathode) with varying flow field patterns and active area of ~ 5 cm2 were produced. Two of the sets with single serpentine flow field of uniform width channels, with 9 passes and 11 passes and a third set of 11 passes with the channel width varying from inlet to outlet (Figure 5). A fourth bipolar plate set with single serpentine, uniform width channels and active area of 5 cm2, manufactured by Fuel Cell Technologies was used for comparative studies.

A fourth bipolar plate set with single serpentine, uniform width channels and active area of 5 cm2, manufactured by Fuel Cell Technologies was used for comparative studies. The three sets were tested in a Scribner Associates Inc., 850C PEM Fuel Cell Test Station with the following Membrane Electrode Assembly (MEA):Electrolyte Membrane: NaFion 117, Anode Catalyst: Pt/Ru/C, Cathode Catalyst: Pt/C, (2.5 mg/cm2 loading).Fig 5.

Bipolar Plate CAD models with varying flow fieldsThe electrochemical performances of the DMFC was evaluated with a single-cell fixture having an active area of 5 cm2 and feeding a preheated methanol solution into the anode at a flow rate of 2.5 mL/min by a peristaltic pump without back pressurization and humidified oxygen into the cathode at a flow rate of 200 mL/min with a back pressure of 20 psi [8].Methanol crossover was evaluated by a voltammetric method [9] in which methanol solution was fed at a flow rate of 2.5 mL/min into the anode side of MEA while the cathode side was kept in an inert humidified N2 atmosphere.

By applying a positive potential at the cathode side, the flux rate of permeating methanol was determined by measuring the steady-state limiting current density resulting from complete electro-oxidation at the membrane/Pt catalyst interface at the cathode side.The performance of the single cells in the fuel cell tests was evaluated by plotting and comparing the scanned potential versus measured current density. As can be seen from figure 5a, the plates produced by indirect SLS worked without fuel leakage and produced expected polarization curves. The plate set with uniform width channels and 11 passes, exhibitedperformance comparable to the commercially manufactured plate set.

The plate set with uniform width channels and 11 passes, exhibitedperformance comparable to the commercially manufactured plate set. The plate set with 11 passes of varying channel width and that with 9 passes of uniform width show diminished performance due to lesser active channel area and higher ohmic and mass transport losses.Methanol crossover tests show that there is a direct correlation to the amount of methanol flowing in the channels to the amount of crossover observed. The larger the active flow area, more the crossover. The plate set with 11 passes of uniform width had the largest amount of methanol crossover, while that with 9 passes of uniform width had the lowest. The flow field of varying channel width showed methanol crossover comparable to that of the commercially produced plates. Since methanol crossover is largely a membrane related phenomenon, these results may be in part due to degradation of the MEA that was used in testing.

Since methanol crossover is largely a membrane related phenomenon, these results may be in part due to degradation of the MEA that was used in testing. Further tests need to be performed using a fresh MEA for each plate set in order to suitably study and compare the effect of the flow field on the amount of crossover observed.(a)Fig 5. (a) Polarization curves of fuel cell test(b) Comparison of methanol crossover4. ConclusionsA process to produce fluid impermeable, electrically conductive graphite bipolar plates by indirect selective laser sintering has been implemented. Bipolar plates have been made using this process and tested in a fuel cell test station and compared with commercially produced plates for performance and methanol crossover. This production and testing process can be further enhanced by the implementation of Computational Fluid Dynamics (CFD) analyses in the design step to optimize the process of flow field pattern selection.

This procedure can be useful in evaluating design concepts aimed at reducing methanol crossover and improving efficiency. Further studies on the effect of heat treatment on the conductivity of the brown part can also be performed. Complete mechanical and thermal characterization of the graphite-cyanoacrylate composite material needs to be performed in order to thoroughly understand the advantages and limitations of infiltrating with ethyl-cyanoacrylate.AcknowledgementThis work was supported by the Office of Naval Research MURI grant No. N00014-07-1-0758.(b)REFERENCES[1] Casalegno, A., Marchesi, R. (2008), DMFC performance and methanol cross-over: Experimental analysis and model validation, Journal of Power Sources, Vol. 185, pp.318–330[2] Liu, H., Songa, C., Zhang, L., Zhang, J., Wang, H., Wilkinson, D. P. (2006), A review of anode catalysis in the direct methanol fuel cell, Journal of Power Sources, Vol. 155, pp.95–110[3] Mehta, V. and Cooper, J.S.

Parts utilizing AlSi10Mg aluminum alloy combine the roles of hydraulic manifold, mechanical interface, and core structure for a meso-scale quadruped robot. These parts allow Naval Research Laboratory (NRL) engineers to design and construct a highly capable quadruped robot that is both dynamic and lightweight. Metal parts were designed with computer aided design (CAD) and constructed using powder-bed direct laser metal sintering (DLMS). High-pressure hydraulic lines with internal fluid passageways were printed into the body assembly seamlessly creating a complex hydraulic manifold. After fabrication, a rigorous program involving post-processing, inspection, and destructive and non-destructive testing was performed to validate the design and manufacturing methods.

After fabrication, a rigorous program involving post-processing, inspection, and destructive and non-destructive testing was performed to validate the design and manufacturing methods. The manifold has been approved for use and integrated on the quadruped, awaiting system level testing.IntroductionThere is current wide-spread enthusiasm for additive manufacturing (AM) technologies with a proliferation of devices leveraging the technique, as well as great anticipation for the vast potential utility and future applications. This is with good reason, as AM removes significant barriers to entry for new manufacturers, including the need for specialized knowledge, reductions in safety hazards, and lower overall cost. The printers and print materials are generally considered safe in significant contrast to traditional subtractive mill and lathe operations that require specialized training of personnel to ensure high quality and safety of operations.

Although plastic AM printing has been prevalent for years now, metal AM printing has recently begun transitioning from primarily an R&D effort to production applications. Although these technologies are less accessible to the general public and significantly more expensive, metal has inherent benefits for high-strength applications. The advent of metal printing opens additional possibilities for AM applications, but also raises questions about whether the product quality is comparable to the high standards of conventional machining techniques. This paper attempts to highlight some of the experiences of researchers at the US Naval Research Laboratory (NRL) in Washington, D.C. as early adopters of metal-based AM technologies.BackgroundNRL conducts a broad range of basic and applied research, and also provides systems engineering, hardware integration and program-level support to government customers for engineering prototypes of advanced systems.

Research projects at NRL are competitively2447awarded based on science and technology merit, and are grouped into focus areas that support the broader objectives of the Naval Science and Technology Strategy. NRL researchers seek out difficult problems for research efforts, with an emphasis on problems of high interest to the military services. Solutions to these problems provide new capabilities, improve soldier health and welfare or streamline operations. As one of these programs, NRL has been investigating advanced hydraulic actuators and controls under a five year 6.2 Discovery and Invention research effort called Meso-scale Robotic Locomotion Initiative (MeRLIn).

The goals of MeRLIn are to develop hydraulic actuators and controls to enable the creation of a highly dynamic quadruped robot that can be built on a scale such that it can be carried in a soldier’s backpack for deployment on demand.The MeRLIn program is inspired by larger payload-carrying quadrupeds build by Boston Dynamics for the Defense Advanced Research Projects Agency (DARPA) over the past decade. Such robots were designed to offload backpacks of deployed soldiers, which can exceed 100 lbs, and overcome terrain too severe for wheeled platforms. Such large payload-capable robots were deployed with Marines for field exercises to demonstrate their capabilities, performance in real- world environments, and potential operational use cases. The conclusions resulting from these exercises were twofold: the robots were too big, and they were too loud.

The conclusions resulting from these exercises were twofold: the robots were too big, and they were too loud. The issue with size was that these complex machines themselves required significant logistical support including specialized handling equipment, spare parts and required significant operational resources including fuel and maintenance. In essence, the large robots were both targets for our adversaries, and required their own supply chain with a burden comparable to their potential payload-carrying benefits. The second concern was noise, which emanates primarily from the internal combustion engine used as the primary power source, and the fear that such a noise source would give away troop positions.

The second concern was noise, which emanates primarily from the internal combustion engine used as the primary power source, and the fear that such a noise source would give away troop positions. Given this feedback, the MeRLIn research team focused on maintaining all the robot’s dynamic capabilities while making it small and portable enough to fit in a soldier’s backpack and operating from rechargeable batteries to address the noise issue.NRL envisions a different use case for near-term robotics, where robots perform covert or hazardous tasks rather than carry significant payloads over long distances. In these use cases, the payload is likely to be a small sensor suite providing situational awareness to the soldier.

In these use cases, the payload is likely to be a small sensor suite providing situational awareness to the soldier. Potential robotic tasks include walking point or flank for a mobile squad, establishing and maintaining a secure autonomous perimeter for an encampment, or performing other hazardous tasks, such as clearing rooms of buildings in a contested environment, transporting supplies between soldiers under fire, performing explosive ordnance operations, or countering the threats of improvised explosive devices (IEDs). These robots are envisioned to have significant dynamic capabilities such that the robot can run faster than a typical soldier, and overcome obstacles by jumping over features such as a culvert, fallen tree, or other debris.

In other applications, the significant dynamic capabilities of the robot can be traded for a less dynamic platform with a larger payload capability in the 5-10 lb range.The key to making a small, dynamically capable robot is maintaining high system-level specific power. High specific power allows the high vehicle-level acceleration necessary for running and jumping, and may enable future capabilities such as climbing. Precursor work at NRL highlighted the inability of small electromechanical actuators to provide sufficient power at these small scales to enable the rapid joint movements required for dynamic operations. Pneumatic systems with pneumatic artificial muscles (PAMs) have also been investigated on separate programs, but a combination of low volumetric efficiency, single-action stroke, and2448complex locomotive controls to overcome the compressibility of the working fluid prevented their use.

As such, the MeRLIn program focused on micro-hydraulic systems to provide the power density needed for this application, and a significant focus of the program has been creating miniature actuator systems including double-acting hydraulic cylinders, legs, pumps, and then integrating them together with sensing systems. A second major thrust of the MeRLIn program is developing and implementing the real-time adaptive control systems necessary to provide responsive and stable operations at the leg and vehicle levels. These controls adapt the effective actuator compliance such that the actuator can be stiff while providing the high force- levels needed for a run or jump maneuver, and then commanding active-compliance to simulate the action of a shock-absorber on the subsequent landing within one cycle of the gait profile.The MeRLIn program uses Additive Manufacturing in several ways.

The most basic is the rapid prototyping of parts to perform fit checks and gain confidence in a parts design before committing more significant resources towards more traditional metal machining vendors for the necessary high-strength parts. One consideration is that operations at the prototype level have to take into account the lower material strength for the temporary plastic parts to ensure that they are used within their capabilities. AM parts have been used routinely as stand-in parts for items with long manufacturing lead times, or as design placeholders during system-level design iterations.The most critical use of AM for the MeRLIn program is to maintain high power density at the system-level. AM contributes to this in two ways, the first of which is that AM allows a single part to be created that acts as a combination of two or more traditionally manufactured parts.

AM contributes to this in two ways, the first of which is that AM allows a single part to be created that acts as a combination of two or more traditionally manufactured parts. As detailed further in the Design section, AM allows for a three part hydraulic manifold to replace multiple fluid system hoses and fittings while functioning as the mechanical structure for the robotic platform. This significantly reduces robot mass and volume by removing individual parts from the system. Additionally, AM provides another function, which is the ability to streamline the mechanical interfaces between parts. This aspect of AM allows the different functions of the robot to be positioned in extremely close proximity to each other and removes significant secondary structures, such as individual mounting interfaces and structural housings.At the onset of the MeRLIn program, the potential benefits of using AM were fairly well established.

Plastic AM parts had become ubiquitous and NRL owned and operated numerous plastic printers, primarily used for creating rapid-design prototypes. Metallic AM parts were just becoming available, either through other government agencies or procurement via commercial vendors. NRL had significant experience with metal AM printing for small mechanical parts, including novel use as fluid manifolds to create small fuel cells for unmanned aerial vehicles.1 The hydraulic manifold discussed in this paper is the most complex and technically challenging metal AM attempted by NRL to date. The manifold design, complete with its complex internal passageways, cannot be built with traditional manufacturing techniques. Additionally, the finished part becomes an integral portion of an active high-pressure hydraulic system and is mated to valves with critically tight sliding fit tolerances that are extremely contamination sensitive.

The use of micron-scale metal powders during the AM process poses a contamination risk to the valve if residuals were to remain in the part, or be released during operation. This requires the manifold to undergo post-processing to purge all remaining powder from the fluid lines. While the use of AM for such a complex part posed numerous challenges, it was viewed as an enabling technology for minimizing robot mass and volume.2449As with any new technology, there were a few known issues to be resolved, some open questions to be answered, and it was recognized that both the technology and implementation had ample room to mature. The MeRLIn program attempted to address these issues in order to take advantage of the significant mass and volume reduction possible using metal AM technologies.

The MeRLIn program attempted to address these issues in order to take advantage of the significant mass and volume reduction possible using metal AM technologies. The designs, applications, processes and findings of these investigations are summarized in the remaining sections of this paper.DesignSystem Design RequirementsThe MeRLIn program builds two hydraulic quadruped robot prototypes during the five year program. The first tethered prototype called ‘Butch’ serves as research platform for a second generation untethered robot prototype named ‘Scout’ that will be built starting in 2018. At the heart of ‘Butch’ is its AM manifold, fed power over a tether from an industrial hydraulic power unit. The design of the manifold was driven by its need to interface with numerous components, support structural loads of the quadruped, efficiently pass hydraulic power to the actuators, and minimize mass.

Although it was desired to make the robot as light as possible, we initially targeted a mass of 22 lb (10 kg) to balance near term manufacturability and dynamic capability. Our target mass was based off of a systems level mass estimate, which was driven by the least scalable component, the hydraulic servo valves. Our 22 lb (10 kg) target falls within the ‘mesoscale’ classification, a driving goal of the MeRLIn program. This target mass was used to size the actuators, which in turn determined the flowrate requirement for the manifold to support. Pictures of the quadruped are shown in Figure 1.Figure 1. Photo and CAD rendering of the hydraulic quadruped, Butch.System pressure was set to the industrial hydraulic standard rating of 3000 psi. Exceeding this pressure would prevent the use of many COTS hydraulic components, both on the quadruped and for the off-board hydraulic power unit.

Exceeding this pressure would prevent the use of many COTS hydraulic components, both on the quadruped and for the off-board hydraulic power unit. At the same time, reducing system pressure would restrict system power density and decrease mobility.Interface and General DesignTo ensure the manifold could interface with all required subsystems, we first decided on driving components. The most significant driving components were the valves, accumulators, and actuators - the valves and accumulators because of their large volume and predetermined2450interfaces, and the actuators because of their restrictive location requirements to maintain leg strength and range of motion. Parts were iteratively located in different configurations until the most viable concept was determined. From there, smaller and more customizable components were added to the system until a detailed design was reached.

From there, smaller and more customizable components were added to the system until a detailed design was reached. During this process, the footprint of the electronics board was selected, driven by both the packaging of Butch, as well as the footprint of the onboard COTS microprocessor. The components the manifold was required to interface with are listed in Table 1.Table 1.

The components the manifold was required to interface with are listed in Table 1.Table 1. Components interfacing with manifoldComponentHydraulic TetherFunctionalityProvides hydraulic powerVendorGoodridgePart description 600 series tubing, 04 sizeMicro accumulatorsProvides high instantaneous hydraulic power in excess of pump flowrateHaweACS/13Subminiature valvesActuator hydraulic linesPressure transducersElectronics boxesInertial Measurement Unit Carrying handles/fall guards Legs Hydraulic actuatorsControls hydraulic actuatorsConnects manifold to actuators Provides system input and return pressure, and up to three DOFs of cylinder pressure Houses NI SOM, as well as all analog and digital I/O.

Custom in-house designUsed for state estimationFor lifting quadruped, as well as protecting from falls For locomotion For moving leg jointsMoogGoodridgeMeggittIn-house (NRL), National Instruments Lord MicrostrainIn-house (NRL)In-house (NRL) In-house (NRL)EO24 (custom two stage flow gain) 600 series tubing, 02 size8510B-2000-DM1N/A3DM-GX4-25N/AN/A N/ADetailed designWe used an iterative routing process to connect the valves to the input pressure and cylinder outputs. The leg mechanical mounts were designed to preserve as much range of motion as practical, and the other interfaces were fit around each other in a puzzle-like fashion. A three- part design was decided on, with two identical front and hind body manifolds, and one central interface manifold. This allowed for a good deal of modularity, ease of assembly, manufacturability, and flexibility to fit within the build volume of commercially accessible metal printers.

This allowed for a good deal of modularity, ease of assembly, manufacturability, and flexibility to fit within the build volume of commercially accessible metal printers. In total, there are 32 discreet fluid lines between the three manifolds as shown in Figure 2 and 3.2451Figure 2. Three-part manifold assemblyFigure 3. Colored supply and return lines, green and blue resp. (left), actuator lines, color indicates one actuator (right)To size the integral fluid interconnects within the manifold, both the flow rate and pressure are taken into account. Based off of the desire for a highly agile system, the valves are designed for a maximum flowrate of 0.7 gpm.

Based off of the desire for a highly agile system, the valves are designed for a maximum flowrate of 0.7 gpm. This flowrate was used to size the cross sectional area of the fluid interconnects using equation2p 15.2LQf  5 d4102where Δp is the pressure drop across the fluid line in psi, ρ is the density of the fluid in lb/ft3, L is the length of the fluid line in feet, Q is the flow rate in gpm, and d is the diameter of the fluid line in inches. The friction factor, f, is given by the Blasius correlation2,2452(1)f316.0Re 14Re 3162Q d where Re is the Reynolds number and μ is the viscosity of the fluid. Combining Equations (2) and (3) into (1). Diameter was chosen not to exceed a pressure drop of 150 psi at maximum flowrate, or 5% of the system pressure.

Diameter was chosen not to exceed a pressure drop of 150 psi at maximum flowrate, or 5% of the system pressure. This resulted in internal passages of 0.06 to 0.12 inches, depending on the location in the manifold.The designed wall thickness of the fluid passages is calculated using the nominal systempressure of 3000 psi, and the Von Mises distortion energy theory.t2 pr i i 2 r r  o i2  1 2 r o 2 r  r2 pr i i 2 r r  o i2   12 r o 2 r    l2 pr i i 2 r r  o i2v   212   3222   132In the above equations, σt is the tangential stress, σr is the radial stress, and σl is the longitudinal stress, correlating to principal stresses σ1, σ2, and σ3, respectively, σv is the Von Mises stress, ri is the tube inner radius, ro is the tube outer radius, pi is the internal pressure, and r is the point at which the stress is measured (ro in this case).

This results in a wall thicknesses ranging from .04” (1.0mm) for the actuators to 0.07” (1.8mm) for the system pressure line. The actuator lines had the lowest factor of safety with a maximum Von Mises stress of 8.2 ksi. This correlates to a 4.1:1 safety factor over the published yield of 34 ksi, and 6.7:1 over ultimate strength.3Because commercially available hose interconnects were mass and volume prohibitive, custom fittings were developed. We developed a miniature conventional plug-in fitting for the hip adduction/abduction and hip flexion/extension joints as shown in Figure 4. For the knee joint port, we added a radial contact ball bearing to allow free rotation of the fitting as the leg moves under pressure. This was implemented due to the large relative motion of the knee actuator, inducing rotation at its port.

This was implemented due to the large relative motion of the knee actuator, inducing rotation at its port. The o-rings of this fitting seal directly against the machined aluminum ports of the manifold, although adding a stainless steel insert for better wear and life was also considered.2453(2)(3)(4)(5)(6)(7)Figure 4. Section view of custom rotary plug-in fittingsThe manifolds were modelled in UG NX 10.0 and exported to vendors as .stl files. The internal piping of the manifolds was modelled using spline paths with boundary conditions matching the ports. Spline parameters were visually adjusted, and measurements were taken to verify wall thickness was preserved between nearby passageways. In parallel, we generated a solid exceeding the dimensions of the manifold with a diamond pattern of lightening voids.

In parallel, we generated a solid exceeding the dimensions of the manifold with a diamond pattern of lightening voids. Then, through a series of combining solid models, performing add/subtract/intersection Boolean operations, and offsetting faces, the diamond pattern replaced much of the volume of the manifold.The only portions of the manifold that were kept solid were the walls of the tubes, a given thickness around any machined features, and several structural ribs and walls to react global loads. This process was especially prone to Boolean errors as offset faces intersected in unstable scenarios. In its most extreme case, the wall thickness of several passageways were left aggressively low, just above a 4:1 burst ratio, because increasing them further caused errors in the model. It is unknown if other CAD packages would have handled this series of Boolean operations with more ease.

It is unknown if other CAD packages would have handled this series of Boolean operations with more ease. Another option was to use a third party program intended for the triangulated files 3D printers use (.STL’s), as these programs give numerous infill options. This option was avoided, as many of the light-weighting features closely reference other features across the manifold. Triangulating the model (i.e. no longer knowing that a cylinder is a cylinder, but instead a surface made of triangles) would have prevented creating many of the precise rib and offset features important for efficient structural loading.Figure 5 shows several steps of the manifold light-weighting process.2454Figure 5. Primary machine features (left), and diamond mesh pattern fill (right).The progression of manifold mass as the design matured and was light-weighted is shown in Table 2. The light-weighted version 2 is the manifold discussed in this paper.Table 2.

Manifold with directly mounted componentsFabricationMaterialAn aluminum alloy (AlSi10Mg) was selected for its low density and high strength to weight ratio. Because of many of the features on the manifold were already at the minimum recommended feature size, going to stronger, denser materials and decreasing volume proportionally was not an option consistent with our minimum mass design.AM ProcessThe first iteration of manifold design was printed out of house by Stratasys Direct.4 This design was used for a single leg testbed, a developmental testbed at NRL that provides hydraulic services and controls for a two or three degree-of-freedom leg. This was a risk reduction effort prior to designing and building the full quadruped. This design was stronger and less aggressively light-weighted than the target design, and weighed 451g. The second iteration24562 2 6 2 8 2 8 12 18manifold and focus of this paper was manufactured by Xometry5, weighing 256g.

The second iteration24562 2 6 2 8 2 8 12 18manifold and focus of this paper was manufactured by Xometry5, weighing 256g. The diamond light-weighting structure and internal manifold ports made this is one of the most complex DLMS parts manufactured by Xometry. Feedback was given by Xometry that we would need to add walls to several areas that would not be self-supporting during the build process; this change was implemented at NRL. An angled build orientation is shown in Figure 7, as selected by Xometry in order to minimize lengths of unsupported structure. Xometry added additional support to all overhangs, but excluded doing so for most of the diamond lattice as the support material would be inaccessible for removal. This, however, incurred risk for the printing defects. For post-processing the gross finished geometry, support material is cut away using a wire EDM, then polished by hand or milled.Build directionFigure 7.

For post-processing the gross finished geometry, support material is cut away using a wire EDM, then polished by hand or milled.Build directionFigure 7. Orientation of the manifold in the build trayIn order to manufacture the manifold, additional material was added at all features that required post machining. These include all threaded features, precision sealing surfaces, and bearing pockets. For all surfaces requiring post machining, .02-.06 inches of material was added, with more material added for larger features.Post processingAll post-process machining was done in-house on a Bridgeport EZ-Trak vertical mill. Features were located by inserting pins into hole features and machining locally. Because the tolerance of DLMS parts linearly increases with distance, features were re-indicated when more than 1-2 inches from each other in order to realign the machine frame with the actual location of the printed features.

It was found that features deviated several hundredths of an inch from their nominal position at these distances. These finding were consistent with published vendor design guidelines, but highlight the fact that maintaining close absolute position tolerances over the span of larger AM parts remains a challenge.In order to reduce frictional flow losses within the manifold, and more so to reduce debris creation during hydraulic operation, the internal passageways of the manifold were extrusion honed. Extrusion honing is a process in which an abrasive slurry (similar to toothpaste in2457consistency) is pushed through a fluid passageway to polish the internal surfaces. To prevent critical surfaces from being abraded during this process, only a select few machining operations to connect the slurry lines were performed prior to extrusion honing.

To prevent critical surfaces from being abraded during this process, only a select few machining operations to connect the slurry lines were performed prior to extrusion honing. Extrusion honing was performed by Extrude Hone AFM.6After receiving the manifold parts from extrusion honing, the remaining machining operations were completed. A rough cleaning was completed with soap and water using an ultrasonic cleaner. Parts were flushed with clean water to remove lose particulates, and then excess water was removed using compressed nitrogen gas. Isopropyl alcohol was flushed through all the ports, blind holes and crevices and then removed with compressed nitrogen gas.

Isopropyl alcohol was flushed through all the ports, blind holes and crevices and then removed with compressed nitrogen gas. Afterwards, the parts were precision cleaned in a Forward Technology vapor degreaser using 3M Novec HFE-7100 solvent in a cleanroom where each port and passageway was flushed with solvent using either port attachment fittings or a small section of 1/16 inch Teflon tubing to get inside each passageway.InspectionUpon coarse visual inspection, a handful of defects and artifacts from the AM process were identified. The main defects were portions of the internal manifold piping with material either caved in or simply missing. An example of this type of defect can be seen below in Figure 8 which looks like a scoop has been taken out of the pipe wall. This discovery raised the question of whether the pipe wall thickness remained constant or the pipe internal cross sectional area was affected.

This discovery raised the question of whether the pipe wall thickness remained constant or the pipe internal cross sectional area was affected. If the wall thickness decreases, then mechanical stress within the wall from the hydraulic pressure increases and creates a stress concentration and structural weakness at these points. If the flow area of the pipe decreases, this would reduce flow rates. We determined that only computerized tomography (CT) scanning technology could investigate these issues nondestructively.Figure 8. Inspection and imaging of manifold pipe defect.In addition to these defects, the main artifacts found appear to be exposed support structure from the build process. An example of these artifacts is shown in Figure 9. Since these artifacts were in difficult-to-reach crevices within the manifold’s geometry, they could not be investigated nondestructively without the use of computerized tomography (CT) scanning technology.2458Figure 9.

Example of artifacts left from the fabrication processCT scanning technology was performed at MQC Labs in Aberdeen, Maryland in order to further investigate these issues non-destructively.7 X-ray CT scanning refers to taking X-ray measurements from multiple orientations to produce cross-sectional or tomographic images. These images, or slices, are then stacked together to create a digital three dimensional model of the part. This process was performed on four body manifolds, two center manifolds, as well as two earlier design versions of the body manifolds.

This process was performed on four body manifolds, two center manifolds, as well as two earlier design versions of the body manifolds. MQC Labs designed the inspection method including how the part was mounted to the rotating pedestal to capture the full part geometry, and used their experience and a number of exposure trials to match the part thickness and density with x-ray exposure to achieve the contrast and resolution needed for close tolerance inspection.The results of the CT scans confirmed that the pipe inside cross sectional area remained constant, and that the wall thickness had significant decreases in the defect areas. This conclusion was made by first inspecting the digital three dimensional composite of the part recreated from the CT scans.

This conclusion was made by first inspecting the digital three dimensional composite of the part recreated from the CT scans. Once the area of interest is defined, the individual CT scanned layer can be extracted from the model and measured at high resolution using the built in measuring tools included in the efX-CT CT analysis software.8 Figure 10 shows an example of each of the two image types. The cross section shown indicates that most of the outside diameter of the pipe is .19” (4.6 mm), the inside diameter is .09” (2.2mm), and the wall thickness is about .05” (1.2mm). At the point of the defect, this wall thickness is .015” (0.4mm), a 60% decrease. The same area was analyzed on the other three body manifolds of the same design, with results detailed in Table 4. It is to be noted that the boundary of the cross section in Figure 10 below is visibly hazy. To account for this, the measurements from the efX-CT software were assumed to be accurate to 0.1 millimeter.2459Figure 10.

To account for this, the measurements from the efX-CT software were assumed to be accurate to 0.1 millimeter.2459Figure 10. 3D composite created from multiple CT scans (left), and an x-ray image of a layer of the cross section (right)Table 4. Measurements of internal piping at defect area.Manifold NumberPart 3 Part 4 Part 5 Part 6Outer Diameter (mm) 4.9 4.6 4.6 4.6Inner Diameter (mm) 2.0 2.1 2.2 2.2Wall Thickness (mm) 1.5 1.2 1.2 1.2Wall Thickness at Defect (mm) 0.6 0.8 0.4 0.5Wall Thickness Decrease 61% 37% 64% 59%At the thinnest section, this would theoretically result in a Von Mises stress of 11.2 ksi, or a factor of safety of 3.0:1 over yield. There is a 5.0:1 FOS over the material’s ultimate stress, resulting in a theoretical burst pressure of 14.7 ksi.

There is a 5.0:1 FOS over the material’s ultimate stress, resulting in a theoretical burst pressure of 14.7 ksi. However, because it is not a uniform cross section, but tapers down to this thickness at one local minimum, the actual burst pressure may be in excess of this.A similar section was anaylzed on a previous design iteration manufactured by Stratasys Direct, as shown below in Figure 11. In this part, there were no major defects present, but the wall thickness is greater than the original part design resulting in a minor reduction in flow area of the pipe. The bulkier mechanical design of the previous design may have had a positive effect to support the piping during printing, or there could have been different AM conditions or processes between the two prints.

Although the differences between these prints is out of the scope of this paper, it would be interesting to have a broader correlation between mechanical design, support material, printer conditions, and the quality of prints for complex build geometries.Figure 11. Example 3D composite (left) and CT layer image (right) without defects.A detailed inspection was also performed of the center manifold parts that connect the two body manifolds together to form the structure of the quadruped as shown in Figure 12. The CT scanning data confirmed that the artifacts in the Xometry parts are exposed support material from the build process. By parsing through the solid layer by layer, a clear grid structure emerges in the void areas of the original part design. We understood that Xometry would need to add2460support structure to the design such that the part could be successfully built, and the geometrical grid pattern in the CT scan confirms its use as shown in Figure 9Figure 12.

3D composite CT scan (left), and x-ray cross sectional layer (right).In terms of model accuracy, the same section detailed in Figure 10 and Table 4 was compared with the engineering model below in Figure 13 and Table 5. The differences between the four manifolds and the engineering model dimensions ranged between 4% to over 30%, or in terms of physical dimensions, from 0.007” to 0.018” difference. As detailed in the fabrication section of this paper, a significant amount of post-machining work was needed for these manifolds; however the internal piping could not be modified or measured due to the complex geometry. Aside from extrusion honing, the internal piping must be largely left untouched. CT scanning data is the only non-destructive method to comprehensively inspect the AM part’s as- built internal geometry relative to the engineering design.Figure 13. CT scan cross section (left), and designed CAD model (right).Table 5. Internal piping dimensions of AM part vs.

CT scan cross section (left), and designed CAD model (right).Table 5. Internal piping dimensions of AM part vs. CAD modelManifold Number CAD ModelOutside Diameter (mm) 4.41Inside Diameter (mm) 2.38Wall Thickness (mm) 1.02Wall Thickness Deviation (%) 0%Part 3 Part 4 Part 5 Part 64.9 4.6 4.6 4.62.0 2.1 2.2 2.21.5 1.2 1.2 1.2+36% +20% +15% +17%2461Burst TestingThe CT scan results of the defect areas for the Xometry parts confirmed pipe wall thickness reductions from the original part design which will result in high stress concentrations at these sections of the piping, and reduced safety margins. To experimentally assess the structural integrity of the manifold assembly, one of the four manifolds parts was pressure tested to failure to confirm the safety of using the AM manifolds at high pressure near humans. The experimental burst test setup is illustrated in Figure 14.

The experimental burst test setup is illustrated in Figure 14. It included a high pressure pump leading to a pressure gauge, then through to an isolated testing burst area where the high pressure line was connected to the manifold with other ports plugged via hoses or threaded bolts to hold pressure. A video camera was set both inside the burst chamber to record the manifold’s failure mode, and outside the chamber facing the pressure gauge to document the precise pressure at failure.Figure 14. Proof test experimental setup. Blue structure indicates secluded burst area.Five separate pressure tests were performed on the same manifold over the course of a single day, fixing observed failure modes along the way. The ideal target pressure was 12,000 psi, four times the maximum normal operating pressure of 3,000 psi, in order to prove safe use in close proximity to humans.

The ideal target pressure was 12,000 psi, four times the maximum normal operating pressure of 3,000 psi, in order to prove safe use in close proximity to humans. Through the five trials, the maximum pressure achieved was 11,000 psi with all failure modes located at the connection points to the manifold or other hardware failures due to the experimental setup, not the manifold itself. A breakdown of the failure modes and maximum pressures held at each trial can be seen below in Table 6.Table 6. Body manifold burst test resultsTest Trial Maximum PressureFailure Mode1 2 3 4 58,600 8,400 9,100 11,000 10,500Hex bolt o-ring Loose hose threading Retaining ring Manifold adapter o-ring Burst hoseThe burst test unit contained the defect that resulted in the minimum pipe wall thickness as determined by visual inspection and later confirmed by CT scanning.

The burst test results are considered the bounding worst case for the lot with the remaining manifolds expected to have higher pressure capabilities. The two manifolds with the thickest walls are currently installed on Butch, with the third reserved as a spare.Figure 15 below shows a frame from the video during burst test trial number five immediately prior to failure. In all tests there was significant leaking above 8,000 psi from the experimental fixturing replacing the valves. This hindered the ability to test past 11,000 without2462a more powerful pump. Despite not reaching the 12,000 psi goal, the results were deemed successful because the manifold held up hydraulically over the course of five tests, including two tests over 10,000 psi. There was one failure not caused by the experimental setup (trial 3), the retaining ring for the rotary plugin fitting sheared through its port at 9,100 psi.

There was one failure not caused by the experimental setup (trial 3), the retaining ring for the rotary plugin fitting sheared through its port at 9,100 psi. However, the geometry of this port allowed modification of the remaining manifolds which had not been post- machined yet, theoretically doubling their strength. The ports of the failed unit were tapped and plugged to allow subsequent testing.Figure 15. Still frame of burst test #5 prior to failure.Fatigue AnalysisFatigue analysis was performed to estimate the fatigue life of the manifold pipe with the defect relative to the nominal wall thickness.

Most of the required material properties had already been experimentally found for sintered AlSi10Mg,10,12 except for the Paris Equation constant, C, for which solid Aluminum 6061-T651 properties were used as a conservative estimate.11 First the calculation was performed using the defect wall thickness of the engineering model with a semicircular initial crack, outlined in Figure 16.9-12Figure 16. Semicircular initial crack geometry in a thick-walled cylinder9Assuming a thin-walled pipe, the hoop stress, σh, is defined as13p hmax         r or or ir i2   2   11     2463(8)where pmax is the maximum fluid pressure.

With inner and outer radii of 0.0469 inches and 0.0626 inches, respectively, and assuming a fluid pressure of 3000 psi, this yields a hoop stress of 10.66 ksi.According to the Paris-Erdogen Law,YKa h (9)where K is the stress concentration, Y is the geometric factor (conservatively estimated to be equal to unity),9 and a is the depth of the crack. Assuming that the value of the stress concentration at failure, KIC, is equal to 36.4 ksi, this yields the crack length at failure, af,KICa Y  h a f f711.3[4.36in []ksiin](10)Since af is greater than the thickness of the pipe at 0.0157 inches, we set the crack length at failure to that. From there, the growth rate of the crack per cycle number, N, is given byda dNKCm YCa m(11)where C is the Paris equation constant for Al 6061-T651, and m is assumed to have a value of 4.

Integrating the crack length, a, from the initial value to the thickness of the tube yields the fatigue life of the component.Nfaai.0CY 015700984 .0 855.1mda m   a 2da  10 66.10 .1 604 610 10cycles 4a 2(12)The same process above was performed for the nominal pipe wall thickness, yielding 58 million cycles required for the pipe to fail. Therefore, nominal to defect wall thickness is a reduction in an order of magnitude; however, 1.8 million cycles is still above the intended scope for this vehicle. If the cyclic pressure was increased from 3,000 psi to 4,500 psi in the worst case thin walled pipe scenario, the amount of cycles would decrease from 1.85 million to 360,000 cycles.ConclusionsThe MeRLIn program created unique high-performance hydraulic manifolds that performed multiple design functions using metal AM techniques.

Although the quadruped has not yet been fully pressurized and actuated, all the individual components have been successfully tested. Despite the flaws that we identified visually and characterized through CT scanning, these parts have been deemed acceptable for use in a research and development environment. Both the2464analytical and experimental burst pressures exceed 3.5:1, but it remains concerning that such a large amount of material is missing. The parts were designed to exceed a 4:1 safety factor over yield stress, and now there is only a theoretical 3:1 margin over yield and 5:1 over burst (for the worst case manifold). However, because a failure in a defect area is expected to be relatively non-violent, easily noticeable, and the area is protected by mounted components, the part will be put into limited use.

However, because a failure in a defect area is expected to be relatively non-violent, easily noticeable, and the area is protected by mounted components, the part will be put into limited use. After fatigue analysis, the defect areas are expected to withstand roughly 1,800,000 cycles, an order of magnitude smaller than the non-defect areas but still above the scope of intended use.The most significant lesson learned was the susceptibility of 3d printing to complex, unsupported geometry. Xometry’s leading hypothesis for the cause of the defect is that the lattice structure resulted in irregularities in heat exposure, causing local hot and cold spots, especially in layers farther from the print bed. When unsupported structure is sintered far away from the supporting lattice, the laser is less effective, resulting in ‘crunchy’ non-structural material which is prone to removal during the cleaning process.

This is compounded by the fact that the powder does not fully sinter during the first pass the laser takes, but finishes sintering over subsequent laser passes. When unsupported geometry is introduced, fresh powder does not have already sintered powder nearby and is prone to failed sintering. However, this is purely speculative and the root cause is still unknown.NRL was mostly out of the loop in deciding where Xometry added support material to the build file. In the future – and in retrospect – we would work more closely with the 3D printing company or operators to identify all areas of concern and add additional support material where appropriate. Had we known the risk of performance-degrading defects, we would have undoubtedly done this. Unfortunately, this would be a very time consuming process due to the large number of unique tubular features across the diamond lattice structure.

Unfortunately, this would be a very time consuming process due to the large number of unique tubular features across the diamond lattice structure. For modeling support material, it would not be recommended to use engineering CAD software, even though the opposite was recommended for generating desired part geometry. This task would be better suited to dedicated 3D printing software that contains all of the printer build constraints.This program answered basic questions about the applicability of metal AM to create challenging high-performance parts. The results of the MeRLIn program help us understand what the AM technologies can and can’t do, where the technology is well suited as well as where it is poorly suited, and where additional margins and constraints should be applied for future similar efforts.

Future manifold builds for the MeRLIn program, including those for the second generation ‘Scout’ robot, will build on the lessons learned through this metal AM experience.AM technology has changed the way we do research at NRL. Scientists use the rapid prototyping capabilities to perform quick design builds to compare alternative approaches or reduce risk that lead to better products or experiments. The rapid turnaround times for manufacturing using these tools allow for broader fields of inquiry than the traditional manufacturing approaches that required detailed design drawings and more expensive parts. These tools allow for faster iteration and more thorough experimentation and results.2465References:1Swider-Lyons, K., “3-D Printing in Hydrogen Fuel Cells for Unmanned Systems”, Spectra, An Official Publication of the Naval Research Laboratory, 2014.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDESIGN, MODELING AND CHARACTERIZATION OF TRIPLY PERIODIC MINIMAL SURFACE HEAT EXCHANGERS WITH ADDITIVE MANUFACTURINGHao Peng1*, Feng Gao2, and Wenjing Hu31. ITAMCO & Atlas3D 2. Dept. of Chemical and Biomolecular Engineering, University of Notre Dame 3. Dept. of Applied and Computational Mathematics and Statistics, University of Notre Dame *Corresponding author: hpeng@itamco.com AbstractNext-generation power plants will generate heated fluids at significantly higher temperatures than current-generation power plants, which challenges the state-of-the-art heat exchanger design. In this study triply periodic minimal surfaces were combined with additive manufacturing for next-generation heat exchanger design.

In this study triply periodic minimal surfaces were combined with additive manufacturing for next-generation heat exchanger design. Triply periodic minimal surfaces separate three-dimensional space into two interpenetrating channels, creating high surface area to volume ratios and low hydrodynamic resistance. Parametric design of triply periodic minimal surface heat exchanger is straightforward because they are governed by simple implicit functions with parameters such as periodic length and offset parameter. In this study a design workflow was developed to streamline the design of triply periodic minimal surface heat exchangers and a numerical model was developed to optimize triply periodic minimal surface heat exchanger design for optimal performance.

Finally, the optimized triply periodic minimal surface heat exchanger was printed with EOS M290 DMLS machine and the performance was tested by experiment.Keywords: triply periodic minimal surface, additive manufacturing, heat exchanger, lattice structureIntroduction1.1 State-of-the-art Heat ExchangersThe most commonly utilized heat exchangers in industrial sectors are shell-and-tube heat exchangers, plate-and-frame heat exchangers, plate-fin heat exchangers and printed circuit heat exchangers [1, 2]. Shell-and-tube heat exchangers consists of a shell with tubes inside it; plate- and-frame heat exchangers consist of a series of corrugated plates supported by a rigid frame; plate- fin heat exchangers consist of side bars, fins, and parting sheets; and printed circuit heat exchangers consist of stacked plates with fine grooves etched into each plate as shown in Fig. 1.2325Figure 1. State-of-the-art heat exchangers.

Thermal effectiveness. The thermal effectiveness is determined by heat transfer coefficients and heat transfer surface area. Fouling. The heat exchanger design should be free from excessive blockage-causing dirt/particulates/other fouling mechanism.The comparison of the state-of-the-art heat exchangers in terms of the above design considerations is shown in Table 1. The shell-and-tube heat exchangers offer the highest maximum temperature and pressure. However, they are usually bulky and heavy due to very low surface area to volume ratio, and not economical for high-temperature and high-pressure applications. Plate- and-frame heat exchangers and plate-fin heat exchangers offer higher surface area to volume ratio than shell-and-tube heat exchangers but they are restricted to low or moderate temperature and pressure applications. Besides, pressure drop in those two heat exchangers is relatively high and they are prone to fouling due to narrow passages.

Besides, pressure drop in those two heat exchangers is relatively high and they are prone to fouling due to narrow passages. Printed circuit heat exchangers achieve high thermal performance and are compact due to the smaller channel geometries. However, they incur larger pressure drops due to long straight micro-channels and they are relatively expensive due to use of chemical etching process [1, 2].A11eatfV (m2 /m3)Tmax (OC)Pressuremax (bar)Shell-and-tube heat exchanger50 - 10011 001000Plate-and-frame heat exchanger120 - 6608 15200Plate-fi n heat exchanger800 - 1500800200Printed circuit heat exchanger200 - 2500980900Table 1. Comparison of different types of heat exchangers in terms of area-to-volume ratio, maximum operating temperature and pressure [1].1.3 AM for Heat Exchanger DesignAdditive manufacturing (AM), also known as 3D printing, emerged nearly three decades ago [4].

Direct metal laser sintering (DMLS) is an AM process for direct generation of metal components [5]. Components are manufactured directly based on CAD files from metal powder feedstock and built up in a layer-by-layer fashion. The metal powders are locally melted by a high- energy laser and reach almost one hundred percent dense microstructure after solidification without2326any additional steps. Thanks to the layer-wise buildup, this process provides components of virtually any shape with almost unlimited freedom of design, and enables manufacturing of complex geometries and internal structures which are not possible with traditional manufacturing techniques.

In addition, AM offers almost unlimited freedom of design, and moreover, a variety of materials such as 17-4 stainless steel, Inconel 625 and 718, aluminum AlSi10Mg and titanium Ti6Al4V can be chosen for different applications.Heat exchangers are a type of components where AM can provide significant advantages in function and efficiency. Figure 2(a) is a multi-furcating heat exchanger designed at General Electric Global Research [1]. The best results showed that the heat exchanger met the pressure drop and heat transfer design requirement with 66% lower weight and 50% lower volume than the conventional heat exchangers. Figure 2(b) shows a heat exchanger with a large number of intricate channels and complex internal structures which are designed to maximize surface area to volume ratio for high efficiency of heat transfer [6]. Figure 2(c) shows a heat exchanger designed and built by EOS® with AM technology.

Figure 2(c) shows a heat exchanger designed and built by EOS® with AM technology. The heat exchanger is self-supporting with integrated cooling fins on outside surfaces [7].(a)Figure 2. (a) Multi-furcating heat exchanger, (b) heat exchanger design with intricate inner channels, and (c) heat exchanger designed by EOS.Triply periodic minimal surface heat exchanger (TPMS-HX) is a new type of heat exchanger which is only possible with AM. The concept of applying TPMS to heat and mass transfer devices was first introduced by Slaughter [8] in 2003. The concept was later expanded by Ryan [9] in 2013. Various types of TPMS were discussed including the Schwartz’s D surface, Schwartz’s P surface, Schoen’s Gyroid Surface and others. However, those studies only stay on conceptual level. Recently Nguyen and et al. [10] presented a study on TPMS-HX used in s-CO2 power cycles. Finite element (FEM) simulation was applied to analyze the flow pattern in a simple Gyroid TPMS structure.

Recently Nguyen and et al. [10] presented a study on TPMS-HX used in s-CO2 power cycles. Finite element (FEM) simulation was applied to analyze the flow pattern in a simple Gyroid TPMS structure. It was reported heat transfer was improved by a factor of 10-100 in comparison to micro-channel heat exchangers. Chandrasekaran [11] performed an experimental study on Schwartz’s D TPMS-HX. The geometry of the TPMS-HX was a simple cube and the material used was Nylon 12. To the best of the authors’ knowledge, existing studies on TPMS-HX are focused on simple cube- or block-type geometries and no metal TPMS-HX is investigated experimentally. This limitation is possibly due to the difficulties of design and manufacturing of complex TPMS-HX geometries by traditional CAD software and manufacturing techniques. In this manuscript, the design, modeling and characterization of an industrial TPMS-HX is presented.

In this manuscript, the design, modeling and characterization of an industrial TPMS-HX is presented. The influence of design parameters of TPMS-HX on the performance of fluid flow and heat transfer is studied by simulation and an EOS M290 powder bed fusion AM machine is employed to build the TPMS-HX for performance testing. The structure of this manuscript is arranged as follows. Section 1 discusses the general design consideration of heat exchanger and the use of AM for TPMS-HX. Section 2 introduces the characteristic of TPMS-HX and the design workflow of2327TPMS-HX. Section 3 presents the simulation of TPMS-HX with COMSOL Multiphysics®. The fluid pattern and heat transfer efficiency for different types of TPMS-HX is analyzed.

Section 3 presents the simulation of TPMS-HX with COMSOL Multiphysics®. The fluid pattern and heat transfer efficiency for different types of TPMS-HX is analyzed. Section 4 shows the experimental studies on the proposed TPMS-HX.Triply Periodic Minimal Surface Heat Exchanger Design2.1 Triply periodic minimal surface (TPMS)TPMS is defined as surfaces of zero mean curvature, meaning that the sum of the principal curvatures at each point is zero. TPMS is a surface minimizes its area with a fixed boundary curve. MS has been mathematically created using level-set equations [12]. These equations consist of implicit trigonometric functions that define an isosurface. Classical examples of TPMSs include Schwartz surface, Gyroid surface and Diamond surface which can be characterized with implicit functions as follows [13].

Figure 3 shows the three types of TPMSs in one periodic length with zero offset parameter.Figure 3. TPMSs in one periodic length: (a) Schwartz surface, (b) Gyroid surface and (c) Diamond surface.2.2 Characteristics of TPMSThe following unique features of TPMSs make them promising in the next-generation heat exchanger design. (1) TPMSs divide a three-dimensional (3D) domain into two separated but interpenetrating channels which provides large surface area to volume ratio. As such, two flows with different temperatures can pass through the separated channels and exchange heat efficiently. Figure 3 shows how cubic domains are separated into Channel A and Channel B by TPMSs. Figure 4 shows that surface area to volume ratio of a Gyroid triply periodic minimal surface increases exponentially with decreasing periodic length and reaches up to 3000 m2/m3 with periodic length of 1 mm.

(4) TPMSs exhibit high structural integrity and can realize pressure equalization [8]. TPMS-HX is usually manufactured by AM as a whole piece without welding or brazing. Also, pressure exerted on two sides of a triply periodic minimal surface tends to equalize. Therefore, they can survive demanding working conditions such as high temperature and pressure. (5) TPMSs are governed by simple implicit functions parametric design of TPMS-HX is straightforward. (6) TPMSs are self-supporting structures and the time-consuming support removal process can be avoided after the AM process.e 3 ~ ;::.2.5 .51 e 2 e " 0 1.5 > g " f " " ~ 0.5dCl)0000000000235 4 Periodic length (mm)6789Figure 4.

Surface area to volume ratio of a Gyroid surface.2.3 TPMS-HX Design ConsiderationsTo achieve optimal heat transfer performance, design parameters of TPMS-HX such as type of triply periodic minimal surface, periodic length, wall thickness, triply periodic minimal surface orientation and offset parameter can be adjusted to accommodate various application scenarios.(cid:120) Type of TPMS: The first step in TPMS-HX design is to determine the proper type of triply periodic minimal surface. There are many types of TPMSs and the most classical ones are Schwartz surfaces, Gyroid surfaces and Diamond surface. Because each type of triply periodic minimal surface gives a unique flow pattern, the heat transfer and pressure drop performance will be also different. For example, the Schwartz surface may give a low pressure drop due to its large straight channels along x, y and z direction.

For example, the Schwartz surface may give a low pressure drop due to its large straight channels along x, y and z direction. However, the heat transfer performance of a Schwartz TPMS-HX may not as good as a Gyroid TPMS- HX because heat transfer in Gyroid surface can be enhanced by its helical-fashion flow pattern. In addition, types of triply periodic minimal surface will influence the manufacturability of TPMS-HX as well as the maximum operating temperature and pressure.(cid:120) Periodic Length and Wall Thickness: The periodic length and wall thickness of TPMS-HX will influence not only the manufacturability but also the heat transfer performance of TPMS-HX. On one hand, the surface area to volume ratio of TPMS-HX increases exponentially with decreasing periodic length as shown by Fig. 4. On the other hand, small periodic length means smaller characteristic dimension which will induce larger pressure2329drop.

4. On the other hand, small periodic length means smaller characteristic dimension which will induce larger pressure2329drop. Small wall thickness is desirable for high heat transfer rates but wall thickness should be larger than a threshold value to guarantee a success build in AM.(cid:120) Orientation of TPMS: The orientation of TPMSs with respect to flow direction is another influential parameter that needs to be determined. There are a few principal orientations for a TPMS along which straight channels can be observed. For example, Fig. 5 shows two principal orientations of a Gyroid triply periodic minimal surface. The first principal orientation is along the edge direction and the second is along the main diagonal direction. Along principal orientations, flow can easily pass through the straight channels with low pressure drop.

Along principal orientations, flow can easily pass through the straight channels with low pressure drop. However, in some cases, principle orientations should be avoided so that flow can exchange heat sufficiently in TPMS.(a)® Flow directionFigure 5. Two principle orientations of Gyroid triply periodic minimal surface structure with respect to flow direction. (a) Flow is along the edge direction. (b) Flow is along the main diagonal direction.(cid:120) Offset Parameter of TPMS: The offset parameter in implicit functions of TPMSs will affect the size of two channel separated by the TPMSs. With a zero offset parameter, TPMSs will divide 3D domain equally into two interpenetrating channels. With a nonzero offset, one channel will expand while the other will shrink. Figure 6 shows the Schwartz, Gyroid and Diamond TPMSs with a positive offset parameter of 0.5. Compared to channels shown in Fig. 3, channel A shrinks and channel B expands.

Figure 6 shows the Schwartz, Gyroid and Diamond TPMSs with a positive offset parameter of 0.5. Compared to channels shown in Fig. 3, channel A shrinks and channel B expands. TPMS-HX with nonzero offset parameter is especially useful when physical properties of working fluids in the two separated channels are significantly different. For example, flow with higher viscosity can be applied to the expanding channel to reduce hydrodynamic resistance and flow with lower thermal conductivity can be applied to the shrinking channel to enhance heat transfer efficiency.Figure 6. TPMSs with positive offset parameter of 0.5: (a) Schwartz Surface, (b) Gyroid Surface and (c) Diamond Surface.23302.4 Developing CAD Models for TPMSSome commercial CAD software can generate simple TPMSs but lots of Boolean operations are needed [13]. An alternative approach is to generate TPMSs directly from their implicit functions.

An alternative approach is to generate TPMSs directly from their implicit functions. In this study, a voxel mesh will be generated for TPMS-HX geometry by evaluating the implicit functions at the center of each voxel. The voxel mesh is converted to STL file by applying the marching cube method. The disadvantage of this approach is that TPMS-HX mesh generated by voxel-based methods is poorly defined at boundaries. This disadvantage can be overcome by reducing the size of the voxel mesh.Fluid Flow and Heat Transfer in TPMS-HX3.1 CFD model for TPMS-HXIn this study, a computational fluid dynamics (CFD) model was developed to predict the fluid flow and heat transfer performance of TPMS-HX. Considering that both fluid flow and heat transfer need to be modeled, a multi-physics CFD simulation software COMSOL Multiphysics® is used. The CAD model of TPMS-HX is imported into COMSOL and the conjugate heat transfer module is selected to model the coupled flow and heat transfer physics.

7(a). Both cold and hot fluids are water and flow in parallel from the inlet end to the outlet end. On the inlet end, the temperatures are 20 °C and 100 °C for the cold and hot fluids, respectively, and fluid velocity is 10 mm/s at the inlets. On the outlet end, outflow boundary is applied for heat transfer physics and zero pressure is applied for the fluid flow physics. Non-slip boundary condition is applied to all boundaries. Insulation boundary condition is applied to the exterior walls and thermally thin approximation is applied to the interior wall (TPMS) with wall thickness of 0.5 mm. Figure 7(b) shows the streamline color coded by temperature in the TPMS- HX. From the inlet or outlet view, a helical-fashion flow pattern can be clearly observed. On one hand, this unique flow pattern help reduce the thickness of the thermal boundary layer on TPMS and enhance heat transfer.

On one hand, this unique flow pattern help reduce the thickness of the thermal boundary layer on TPMS and enhance heat transfer. On the other hand, the helical-fashion flow prevents dirt or particulates from attaching to the TPMS and hence extends the serving life of the TPMS-HX.Figure 8 shows the evolution of velocity in the first two periodic lengths of the TPMS-HX. The velocity pattern is taken from the cross-section of yz plane along the x axis when the flow reaches steady state. The flow velocity is significantly higher along the straight channels in the principle orientation as shown in Fig. 5(a). The velocity at the center of the straight channel is up to 50 mm/s, five times of the inlet velocity. The flow away from the center of the straight channel follows the surface gradient of the TPMS, spins about the center of the straight channel as moving forward, and eventually forms the helical-fashion flow.

In addition, the velocity distribution is eccentric about the center of the straight channel as highlighted in Fig. 8, and the eccentric direction changes in counter-clockwise direction for this TPMS-HX design. By comparing velocity pattern at the same position of the first and second TPMS unit, one can observe that the flow is fully developed in the first periodic length.First period lengthmm/s:::: II "4035II"30§25Eccentri velocity distribut'V)II"2015II"10§50 N II"0Figure 8. Evolution of flow velocity in the first two periodic length of the TPMS-HX.23323.3 Heat Transfer in TPMS-HXFigure 9 shows the temperature evolution in the TPMS-HX. The hot flow is 100 °C and the cold flow is 20 °C at the inlet. After one periodic length (x=10 mm), the temperature of the cold and hot flow changes significantly, especially the flow away from the center of the straight channels.

After one periodic length (x=10 mm), the temperature of the cold and hot flow changes significantly, especially the flow away from the center of the straight channels. At the outlet (x=100 mm), the mixing-cup temperature of the hot and cold temperatures are 62.3 °C and 57.8 °C, respectively. The mixing-cup temperature is calculated with(cid:179) (cid:179)u n (cid:152)ds(cid:85)C T pDT(cid:32)mixing cup(cid:16)dsC (cid:85)u n (cid:152)pD where D is the surface area of the flow at outlet, u is flow velocity at outlet and n is the normal direction on outlet surface.- - - x= 0x=3 0mmx= 60mmx=20mmx = 50 mmx = 100 mmFigure 9. Evolution of temperature in the TPMS-HX.To show the efficiency of the TPMS-HX, the TPMS-HX is benchmarked against a traditional plate heat exchanger with straight channels. The dimensions of the traditional heat exchanger are the same as the TPMS-HX.

The dimensions of the traditional heat exchanger are the same as the TPMS-HX. Two hot flow domains are sandwiched with two cold flow domains to resemble the flow condition in the TPMS-HX. Figure 10 show the comparison of those two type of heat exchanger at inlet velocity of 10 mm/s and 20 mm/s. The interfacial surface area between the cold and hot flow is 251.5 cm2 in TPMS-HX and 120 cm2 in the reference heat exchanger as shown in Fig. 10(a). At the outlet, the mixing-cup temperature is significantly different for those two types of heat exchangers as shown in Fig. 10(b). The hot and cold fluid temperatures in TPMS-HX are much closer, showing a difference of 4.5 °C for inlet velocity of 10 mm/s and 7.9 °C for inlet velocity of 20 mm/s, as opposed to a difference of 42.4 °C for inlet velocity of 10 mm/s and 52.5 °C for inlet velocity of 20 mm/s in the reference heat exchanger. This is due the high heat transfer rate in TPMS-HX. Figure 10(c) shows the maximum pressure drop.

This is due the high heat transfer rate in TPMS-HX. Figure 10(c) shows the maximum pressure drop. The pressure drop in TPMS-HX is also much higher than that in the reference heat exchanger. Maximum pressure drop in TPMS-HX with inlet velocity of 20 mm/s is 19.15 Pa while the counterpart in the reference heat exchanger is only 1 Pa. Figure 10(d) and 10(e) shows the comparison of heat transfer rate and heat transfer coefficients. Heat transfer coefficients are calculated as2333degC1009080706050403020(5)P (cid:16) where P is the total exchanger power and A is the surface area through which P flows through. The heat transfer rates in the TPMS-HX are 630.9 W with inlet velocity of 10 mm/s and 1189.2 W with inlet velocity of 20 mm/s, 5.4 times and 7.5 times of their counterpart in the reference heat exchanger respectively.

Comparison of the performance of TPMS-HX to a traditional plate heat exchanger.Manufacturing of TPMS-HX with AM4.1 TPMS-HX with AM Figure 11 shows a TPMS-HX design and the cut-view of the design. The core of the TPMS-HX is a gyroid TPMS with wall thickness of 1 mm. In addition, boundary walls, inlets and outlets needs to be added to the TPMS to properly separate the hot and cold fluid. The TPMS-HX design was then imported to Atlas3D® (https://atlas3d.xyz/) to find a proper build orientation and generate necessary support structures as shown in Fig. 12(a). Atlas3D® is a cloud-based platform for optimizing build orientation and generating support structures for parts in AM. The oriented STL and support files were directly downloaded from Atlas3D® and imported to EOSPRINT. EOS M290 machine was employed to fabricate the TPMS-HX. The laser power is 220.1 W, laser scanning speed is 755.5 mm/s and the hatch distance is 110 microns.

EOS M290 machine was employed to fabricate the TPMS-HX. The laser power is 220.1 W, laser scanning speed is 755.5 mm/s and the hatch distance is 110 microns. Metal powder feedstock 17- 4PH stainless steel was used and 40 micron build layer was chosen. Figure 12(b) shows the final product of the TPMS-HX.2334Figure 11. (a) TPMS-HX design. (b) Cut-view of the TPMS-HX design.Figure 12 (a) TPMS-HX and its support structures in Atlas3D viewer. (b) TPMS-HX printed with EOS M290 machine.23354.2 Manufacturability of TPMS-HX by AMFigure 13 shows a TPMS-HX with periodic length of 10 mm and wall thickness of 0.2 mm printed with a layer thickness of 40 microns in an EOS M290 machine. The TPMS-HX failed in leak test and after cutting the TPMS-HX with electrical discharge machining (EDM), a uniform distribution of voids on the TPMS can be observed. Those voids were probably due lack of fusion and were avoided by increasing the wall thickness to 1 mm in the new design.

Those voids were probably due lack of fusion and were avoided by increasing the wall thickness to 1 mm in the new design. Influential factors on manufacturability of TPMSs include the types of triply periodic minimal surface, periodic length, wall thickness, build orientation and etc. The periodic length of the TPMS should be within a threshold so that heat at overhanging regions of TPMSs can dissipate efficiently. Small wall thickness is desired to reduce thermal resistance. However, strength and integrity of TPMSs also decreases with decreasing wall thickness.Figure 13. A TPMS-HX failed on leaking test due to voids on the TPMS.ConclusionIn this study we proposed a novel heat exchanger design by utilizing the TPMS. The performance of the heat exchanger was demonstrated by developing a CFD model in COMSOL. Simulation results showed that heat transfer rates were improved by a factor of 7.5 times in comparison to a traditional plate heat exchanger.

Simulation results showed that heat transfer rates were improved by a factor of 7.5 times in comparison to a traditional plate heat exchanger. Finally, an industrial-size TPMS-HX was printed with an EOS M290 DMLS machine and manufacturability study on the TPMS-HX showed that a wall thickness larger than 0.2 mm was desired.1.2.3.4.References Ohadi, M., et al., Recent Developments in High Temperature Heat Exchangers: A Review. Frontiers in Heat and Mass Transfer, 2018. 11. Li, Q., et al., Compact heat exchangers: A review and future applications for a new generation of high temperature solar receivers. Renewable and Sustainable Energy Reviews, 2011. 15(9): p. 4855-4875. Southall, D., R.L. Pierres, and S.J. Dewson. Design Considerations for Compact Heat Exchangers. in Proceedings of the 2008 International Congress on Advances in Nuclear Power Plants 2008. Anaheim, CA USA. Gibson, I., D.W. Rosen, and B.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDESIGN, SIMULATION AND EXPERIMENTAL INVESTIGATION OF 3D PRINTED MECHANICAL METAMATERIALSSunil Magadum†*, Amol Gilorkar*, Deepak M*, Rakshith B S*, Navaharsha P*, Nagahanumaiah† and Somashekara M A*†Central Manufacturing Technology Institute, Bengaluru 560022 *Department of Mechanical Engineering, Indian Institute of Technology, Dharwad 580011AbstractMechanical metamaterials have generated special interest recently due to their tailorable structure, exceptional mechanical properties, and advancements in 3D printing processes that allow the fabrication of intricately structured components. Designing innovative structures of metamaterials will lead to the development of advanced materials with special properties.

Designing innovative structures of metamaterials will lead to the development of advanced materials with special properties. The experimental investigation presented in this paper involves the design, simulation, fabrication, and testing of three different mechanical metamaterial models i.e. Chiral, Re-entrant, and Hybrid printed in acrylonitrile styrene acrylate (ASA) using fused deposition modeling (FDM). Subsequently, a uniaxial compression test and ex-situ characterization was performed for studying the mechanical properties, the types of fracture and crack propagation of the printed metamaterial models which may lead to the development of metamaterials with tunable compressive/bending stiffness.Keywords: 3D Printing, Mechanical Metamaterials, Chiral, Re-entrant, Hybrid Model1. IntroductionA metamaterial is a non-natural architecture design that is produced to get extraordinary properties that are not present in naturally existing materials [1].

IntroductionA metamaterial is a non-natural architecture design that is produced to get extraordinary properties that are not present in naturally existing materials [1]. Repeated unit patterns of small to large scales are used to get the desired metamaterial design. The properties of the materials such as stiffness, poisson’s ratio, force absorption capability, or thermal expansion can be manipulated using internal structures. The external properties of the materials also can be enhanced by using optimized parameters of the unit cells such as the shape of the unit cell, size and unit cell orientation, etc. Hence by changing the internal geometry of the material we can control its properties. Metamaterials research has involved broader areas such as optics, communication, biomedical, and so on.

Hence by changing the internal geometry of the material we can control its properties. Metamaterials research has involved broader areas such as optics, communication, biomedical, and so on. Metamaterials are generally classified as electromagnetic, mechanical, acoustic, and thermal metamaterials [2-5].Mechanical metamaterials are man-made materials that derive their unusual properties from the geometry of building blocks rather than the properties of their constituents [6]. The derived unusual properties of these metamaterials include negative rigidity, negative poisson’s ratio, negative squeezability, and fading shear modulus [7]. Mechanical metamaterials are further classified as Auxetic, Penta-Mode, and Origami-Based Metamaterials.2Fig. 1: Mechanical metamaterials classificationThe auxetic metamaterials are associated with a negative poisson’s ratio and become thicker normal to the applied force when stretched.

1: Mechanical metamaterials classificationThe auxetic metamaterials are associated with a negative poisson’s ratio and become thicker normal to the applied force when stretched. This is possible only due to their internal structure and the way they deform under uniaxial loading. Gibson et al invented the first conventional re-entrant Metamaterial structure in 1982 [8]. Apart from re-entrant various other types of auxetic Metamaterial have been developed such as chiral, rotating polygon, etc. as explained by Liu Y et al. [9]. According to the authors Parth Uday Kelkar, et al. [10] materials have different poisson’s ratio depending upon their internal geometry. Auxetic metamaterials are used in bio-medicine, soft robotics, and acoustics. As per Kolken, et al. [11] these materials also find applications in the areas of automotive, aerospace, and defense industries due to their improved shock absorbance properties.

As per Kolken, et al. [11] these materials also find applications in the areas of automotive, aerospace, and defense industries due to their improved shock absorbance properties. Based on the unit cells used in printing, various types of auxetic metamaterials were discussed by Xin Wu et al. [12]. Another class of mechanical metamaterials is Penta-mode Metamaterial which is classified based on bulk/shear modulus properties. Normally Penta-mode metamaterials have a larger bulk modulus in comparison with their shear modulus and have a poisson’s ratio value of 0.5[13-16]. On the other hand, the Origami metamaterials which are derived from Japanese paper folding art find their applications in bio-medical stents and self-folding robotic devices [17-19].As the auxetic mechanical metamaterials are finding more significance in recent times, there is a huge demand for the development of new metamaterial models based on their auxeticity to cater to a wider spectrum of applications.

To explore and develop new mechanical metamaterials, it has been decided to conduct an experimental investigation on 3D printed mechanical metamaterials and study their mechanical properties. In the current work presented in this paper an attempt is made to experimentally investigate the two established auxetic mechanical metamaterial models namely chiral and re-entrant. Also based on these models a hybrid metamaterial model (Anti-tetra chiral) was developed and tested under a uniaxial compression test to study its mechanical properties and compare it with the well-established chiral and re-entrant models.2. Experimentation2.1 3D Printing of Mechanical Metamaterial Models All three models chiral, re-entrant, and hybrid as shown in Figure 2 were fabricated using Stratasys F120TM 3D-printer.

Experimentation2.1 3D Printing of Mechanical Metamaterial Models All three models chiral, re-entrant, and hybrid as shown in Figure 2 were fabricated using Stratasys F120TM 3D-printer. The parameters like bed temperature, nozzle temperature, printing speed, infill density were all held to constant values as these are the default settings in Stratasys F120TM 3D- printer according to the material selected. The printed models were cured before mechanical testing by keeping them at room temperature for about 24 hours.3Fig. 2: 3D Printed Mechanical Metamaterials (a) Chiral, (b) Re-entrant, and (c) Hybrid ModelIn this study, Acrylonitrile Styrene Acrylate (ASA) filament was used to fabricate the samples using the FDM process. ASA material is selected over ABS (acrylonitrile butadiene styrene) which is normally used in recent FDM fabrication as it displays better mechanical properties compared to ABS material.

ASA material is selected over ABS (acrylonitrile butadiene styrene) which is normally used in recent FDM fabrication as it displays better mechanical properties compared to ABS material. The properties of ASA material are tabulated below in Table 1.Table 1: Properties of ASAProperties Glass Transition Temp. Density Compressive Strength Young’s ModulusTypical values 112 0C 1.08 g/cm3 75 MPa 2.5 GPaThe Chiral, Re-entrant, and Hybrid Metamaterial models (Fig. 2) were fabricated by varying printing layer height (0.178mm and 0.254mm) with a constant rib thickness of 0.75 mm.

2) were fabricated by varying printing layer height (0.178mm and 0.254mm) with a constant rib thickness of 0.75 mm. The other dimensions of the models are given in Table 2.Table 2: Dimensions of Metamaterial modelsModel Chiral Re-entrant HybridLength, l 50.59 50.76 49Width, w 50.59 47.13 49Height, d 20.80 20.00 20.002.2 Uniaxial Compression Testing of Printed Models To investigate the mechanical behavior of all three metamaterials, a uniaxial compression test was conducted using BISS UTM with 100 kN load carrying capacity. The models were tested under two different strain rates of 0.5 % (5×10-3 s-1) and 1 % (10×10-3 s-1). Figure 3 shows the deformation of 3D printed metamaterials under the uniaxial compression test.4Fig. 3: Deformation of printed metamaterials under uniaxial compression test with 0.5% s-1 strain rate2.3 Simulation Study:The deformations of the designed models were studied through quasi-static analysis using ANSYS software.

The deformation modes of the models is given in Figure 4.Fig. 4: Deformation modes of models2.4 Ex-situ Characterization:5The printed models after undergoing uniaxial compression tests are characterized using scanning electron microscopy (SEM) to study the crack and fracture morphology of the models. Different fracture and crack propagations were observed and are discussed in the next section.Abbreviations used: C0.178_0.75mm_0.5% [R stands for Re-entrant and H stands for Hybrid Model] Chiral with 0.178 mm layer height and 0.75mm rib thickness tested under 0.5% strain rate100 µm100 µmChiral Before Test : C0.178_0.75mm_0.5%After Test : Crack depth ratio is 0.964100 µm200 µmChiral Before Test : C0.254_0.75mm_0.5%After Test : Crack depth ratio is 1.128(a).

6: Fracture morphology of the models3. Results and DiscussionA uniaxial compression test was conducted on all three metamaterial models as shown in Figure 3 to compare the mechanical behavior of the models and to investigate the effect of layer height on their mechanical properties.The deformation modes under uniaxial compression tests were modeled using ANSYS software as shown in Figure 4 & compared with experimental results. The images of the cracks developed after testing as shown in Figure 5 prove that, progressive crack propagation is observed in both the chiral and re-entrant models. Whereas in the case of the hybrid model the failure is due to brittle fracture. Crack depth ratio (length of crack/rib thickness) of 0.964 & 1.128 is observed in the case of the chiral model printed with a layer height of 0.178 mm & 0.254 mm respectively.

Crack depth ratio (length of crack/rib thickness) of 0.964 & 1.128 is observed in the case of the chiral model printed with a layer height of 0.178 mm & 0.254 mm respectively. Whereas re-entrant model has shown the crack depth ratio of 0.928 mm & 1.292 mm for layer height of 0.178 mm & 0.254 mm respectively.Figure 6 shows the fracture morphology of models after the uniaxial compression test. In the chiral model, a combination of inter and intra-laminar fracture was observed for 0.178 mm layer height whereas for 0.254 mm layer height only intra-laminar fracture was observed leading to the uniform fractured surface. A similar trend is also observed in the Re-entrant model. But for the hybrid model in both layer height cases, a combination of inter and intra-laminar fracture were observed. This may be attributed to the brittle fracture that occurred during testing.8(a) Chiral(b) Re-entrant(c) HybridFig. 7.

This may be attributed to the brittle fracture that occurred during testing.8(a) Chiral(b) Re-entrant(c) HybridFig. 7. Effect of layer height used for printing on the true stress-strain relationship under strain rate of 0.5% s-1 and 1% s-1 respectively.9(a)(b)Fig. 8. True stress-strain relationship between hybrid, re-entrant, and chiral models printed using a layer height of (a) 0.178 mm and (b) 0.254 mm under strain rate of 0.5% s-1 and 1% s-1 respectively.The effect of layer height used in the printing of the models is studied and is shown in Figure 7. As it is evident from the above shown figures that an increase in the strain rate from 0.5 % to 1 % has no significant effect on the mechanical properties. With an increase in layer height of around 30 %, there is an increase in peak stress of 16 %, 25 %, and 40 % for chiral, re-entrant, and hybrid models respectively.

With an increase in layer height of around 30 %, there is an increase in peak stress of 16 %, 25 %, and 40 % for chiral, re-entrant, and hybrid models respectively. This can be attributed to the reduction of interfaces between the layers.For the chiral model fabricated with a layer height of 0.178 mm as shown in Figure 8 (a), during a uniaxial compression test, the stress increases to a peak value of 0.6 MPa at around 10 % of strain and then maintains more or less this value till 35 % strain due to the self-contact between the ribs. Hardening of the chiral structure occurs beyond 35 % strain. Whereas in re-entrant model, peak stress of 1.2 MPa was observed at 5 % strain followed by a sudden drop till 10 % strain and maintaining it with minor softening. Once the strain crosses ~45%, the hardening takes place.

Once the strain crosses ~45%, the hardening takes place. And for hybrid (Anti-tetra chiral) model, the stress increases linearly to a peak value of around 1.8 MPa at 10 % strain and then slowly reduces and reaches almost half the peak value at 35 % strain and hardens. The hardening of structures occurs at 35 % strain for both chiral and hybrid models, whereas for re-entrant the hardening occurs at around 45 % strain. A similar trend has been observed for the models printed with a layer height of 0.254 mm as shown in Figure 8 (b). The sudden drop in the stress value 10in the case of re-entrant may be attributed to the reduction in auxetic effect due to its instability under small deformation (~4%). Whereas in the case of the chiral model, the hardening occurs earlier compared to re-entrant due to drastic reduction in its volume under test and it conserves the auxetic effects as reported by Yunyao Jiang et al.

[20].It is evident from Figure 8 that the hybrid model (Anti-tetra chiral) has shown better stiffness compared to both chiral and re-entrant models and has shown an improved chirality as compared to the re-entrant model. This can be attributed to the anti-chirality nature of the structure.4. ConclusionThe fused deposition modeling (FDM) process was used for the fabrication of three different mechanical metamaterial models namely chiral, re-entrant, and hybrid (Anti-tetra chiral). The layer height effect on the mechanical properties of all three models were studied and observed that the increase in layer height increases the strength of the models. Also, a hybrid model which is a combination of both chiral and re-entrant models was developed successfully and tested. It is observed from the results of the experimentation that the hybrid model has shown better mechanical properties as compared to chiral and re-entrant models.

It is observed from the results of the experimentation that the hybrid model has shown better mechanical properties as compared to chiral and re-entrant models. This shows that the improved properties can be achieved by changing and combining the properties through its structure depending upon the required application. Also in this, for both chiral and re-entrant models a combination of inter and intra-laminar fracture was observed for 0.178 mm layer height. Whereas for 0.254 mm layer height only intra-laminar fracture was observed leading to the uniform fractured surface. And in hybrid model a combination of inter and intra-laminar fracture were observed.5. AcknowledgementsThis work is supported by the Science and Engineering Research Board, Department of Science & Technology, Government of India, under Start-up Research Grant (SRG) scheme (Project no: SRG/2019/000990).

An oxide-carbide-nitride system consisting of ceramic powders and pre-ceramic organics, designed to yield ZrB2 after reaction sintering, has been developed to produce an aqueous-based extrudate for subsequent processing in the CODE system. Pressurelessly sintered test specimens containing 1 wt% PVA binder achieve high relative density ≥ 99%. The viscoelastic response of the extrudate was characterized via spindle rheometry with a small sample adapter. Batches with 1 wt% PVA and 0.5 wt% Methocel show strong shear thinning characteristic, under shear rates of 1-28 s-1. XRD and SEM were utilized for microstructural analysis to determine phase development and microstructural morphology.1. IntroductionZirconium diboride (ZrB2) is classified as an ultrahigh temperature ceramic (UHTC), all of which have melting temperatures over 3000oC [1]. This class of material performs well in harsh environments, often exceeding the capabilities of traditional engineering materials.

This class of material performs well in harsh environments, often exceeding the capabilities of traditional engineering materials. ZrB2 has a high melting temperature (3245oC), high Young's modulus (526 GPa), high hardness (23 GPa), and low theoretical density (6.09 g/cm3) [2, 3, 4]. The application of additive manufacturing for UHTCs may provide a cost-effective method for fabrication of complex components for next-generation gas turbines, rocket engines, and hypersonic vehicles [5, 6].Sintering of non-oxide ceramics is often difficult because of their high melting temperature, due to strong covalent bonds (metallic, ionic, and covalent in the case of ZrB2). Sintering to high densities is normally achieved through solid-state sintering by pressure-assisted methods, such as hot pressing and hot isostatic pressing.

Sintering to high densities is normally achieved through solid-state sintering by pressure-assisted methods, such as hot pressing and hot isostatic pressing. However, these methods limit production to the creation of monolithic geometries, and as such, are not viable for complex geometries often associated with additive manufacturing. Alternative pressureless processing routes have proven equally successful [7, 8]. Pressureless sintering is the promotion of densification without the application of an external pressure upon a body; this is the most common sintering method and often what is referred to when ‘sintering’ a ceramic body [9].

Multiple sintering additives have been found to assist in pressureless sintering of ZrB2: 1.7 wt% carbon sintered at 1900oC and 4.5 wt% B4C sintered at 1850oC resulted in near theoretical density for powder attrition milled with tungsten carbide (WC) media [10].929Reaction sintering is a method in which both chemical reaction and densification occur in a single sintering cycle. Normally a method reserved for production of high purity powders, several reaction paths have been introduced with varying thermodynamic requirements [11, 12, 13]. However, none have been applied to in situ fabrication using additive manufacturing (AM). With application of colloidal processing techniques, an extrudate may easily be designed, such that it has the desired rheological properties for deposition, the chemistry for reaction sintering, and a particle size to promote specific phase development [14, 15].

These parameters can be further controlled to promote high relative density without pressure-assisted sintering, such that complex, near net-shape components may be fabricated. Elimination of the machining costs associated with ceramic materials would alone justify in situ fabrication. The structural and thermal properties of ZrB2 are valuable to many industries that would favor a method of production that provides a high degree of compositional control with low machining costs.Ceramic On-Demand Extrusion (CODE) is an extrusion-based AM process which can successfully fabricate mesoscale artifacts. Using high solids loading (>50 vol%), aqueous, ceramic extrudate, deposition is done at room temperature to build a 3D geometry layer-wise. After deposition, each layer is partially solidified by uniform infrared radiation applied perpendicular to the top face.

After deposition, each layer is partially solidified by uniform infrared radiation applied perpendicular to the top face. Concurrently, the built portion is surrounded, flush with the top most layer, by a low molecular weight oil to promote one-dimensional drying. Dynamic drying eliminates a water content gradient in the deposited part, imparts greater stability during the build, and enables the production of fracture and warpage free ceramic parts [16].Pressureless reaction sintering, applied to CODE, provides a solution to in situ fabrication of near net-shape UHTC components with a high degree of microstructural control and minimal post-processing that has not yet been achieved in AM to date. This study is the initial investigation to design an oxide-carbide-nitride extrudate to fabricate highly complex near net-shape components.

This study is the initial investigation to design an oxide-carbide-nitride extrudate to fabricate highly complex near net-shape components. The goals of this study are as follows: 1) identify viable reaction materials, 2) determine the ratio of additives to promote pressureless sintering to achieve high relative density, 3) maintain nominal purity with high relative density, 4) achieve a high degree of microstructural control, and 5) compare the mechanical properties between pressureless and reaction sintered ZrB2 parts fabricated by Ceramic On-Demand Extrusion.2. Experimental ProcedureCommercially available zirconium diboride (ZrB2, Grade B, H.C. Starck, Karlsruhe, Baden-Wüttemberg, Germany) and boron carbide (B4C, Grade HS, H.C. Starck, Karlsruhe, Baden- Wüttemberg, Germany) powder were used in this study for pressureless sintering trials to establish a baseline sintering temperature.

Starck, Karlsruhe, Baden- Wüttemberg, Germany) powder were used in this study for pressureless sintering trials to establish a baseline sintering temperature. The reported average particle size was in a range of 1.5-3.0 µm with a 97% (metal basis) purity for ZrB2; the average particle size of the B4C powder was 0.8 μm, with a surface area of 15-20 m2/g, and a B:C ratio of 3.7-3.8.To prepare an extrudate for in situ fabrication of ZrB2, zirconia (ZrO2; Inframet Advanced Materials, Manchester, Connecticut, USA), boron nitride (BN; Grade F 15, H.C. Starck, Karlsruhe, Baden-Wüttemberg, Germany), and graphite (Carbon Black, Cabot, Alpharetta, Georgia, USA) were attrition-milled in methyl ethyl ketone (MEK, Sigma Aldrich, St. Louis, Missouri, USA) at 600 rpm for 2 h; BN had a reported average particle size of 6.0 μm and surface area of 10-20 m2/g.930A high molecular weight polyvinyl alcohol (PVA; avg. molecular weight 70,000-100,000, Sigma Aldrich, St.

molecular weight 70,000-100,000, Sigma Aldrich, St. Louis, Missouri, USA) and cold water dispersible hydroxypropyl methylcellulose (Methocel J5MS, DOW, Midland, Michigan, USA) solutions were used as water-soluble binders.The surface area of as-received powder was measured using nitrogen adsorption analysis according to Brunauer-Emmett-Teller (Nova2000e, Quantachrome Instruments, Boynton Beach, Florida, USA) and the particle size distribution by laser diffraction (S3500, Microtrac, York, Pennslyvania, USA). Sedimentation studies were used to approximate the best concentration of dispersant on a surface area basis, using Darvan C-N (ammonium polymethacrylate; Vanderbilt Minerals, LLC, Norwalk, Virginia, USA) and Dolapix CE 64 (carbonic acid; Zschimmer & Schwarz, Lahnstein, Rhineland-Palantinate, Germany).

In the known pH regime for stability, 0.5- 2.0 mg/m2 dispersant was added to respective 10 mL graduated cylinders; 3 g of powder was added to each cylinder; 10 mL pH adjusted solvent was added; each cylinder was covered using paraffin film, agitated, then ultrasonicated for 600 s before being left stationary for sedimentation. The pH value was measured using a pH meter (HI 2210, Hannah Instruments, Woonsocket, Rhode Island, USA) and adjusted dropwise using an ammonium hydroxide solution (NH4OH, 30% NH3 basis, Sigma Aldrich, St. Louis, Missouri, USA).2.1 Pressureless SinteringA ZrB2 composition with 4 wt% B4C additive was used for pressureless sintering studies. These powders were attrition-milled in a polymer-lined jar using cobalt-bonded tungsten carbide (WC) media at 600 rpm for 2 h in acetone or MEK. After milling, the powder was retrieved via rotary evaporation (Rotovapor R-124, Bucchi, Flawil, St. Gallen, Switzerland).

After milling, the powder was retrieved via rotary evaporation (Rotovapor R-124, Bucchi, Flawil, St. Gallen, Switzerland). The powder was then crushed using a diamonite mortar and pestle and passed through a sieve stack: 250 µm / 60 mesh; 150 µm / 100 mesh; 106 µm / 140 mesh; and 90 µm / 170 mesh. Collected powder was uniaxially pressed into pellets for a sintering study. Two samples, each of five pellets, were made using a laboratory benchtop press (model 3851-O, Carver, Wabash, Indiana, USA); one set with and the other without PVA-binder. Due to incompatibility between PVA and MEK, batches that included binder were milled in acetone. After the doped ZrB2 powder was retrieved, PVA binder was introduced as a 10 wt% solution, and the mixture was milled in a HDPE wide-mouth bottle with WC media for 2 h to homogenize. Solvent was removed by boiling the slurry under a high (700-1200) rpm stir until dry.

Solvent was removed by boiling the slurry under a high (700-1200) rpm stir until dry. For each pellet, 1.5 g of powder was loaded into a 12.7 mm die and pressed for 60 s at 70.3 MPa; between compaction, the die was cleaned with acetone and lubricated with stearic acid.The samples were sintered individually; a 6.35 cm dia. x 5.33 cm tall cylindrical graphite crucible was lined with flexible graphite foil (GRAFOIL, GrafTech Intl., Parma, Ohio, USA) and coated with an aerosol boron nitride lubricant (BN SP-2018, Materion, Buffalo, New York, USA) for sintering. Once dry, the crucible was loaded and topped with a permeable graphite foil lid. The sintering schedule was carried out in a resistance-heated, graphite, bottom-loading furnace (model 1000-4360-FP30, Thermal Technology Inc., Santa Rosa, CA, USA). The furnace was ramped from room temperature to 700oC at 10oC/min in flowing argon. Once at temperature, the furnace was held for 0.5 h before ramping to 1450oC at 10oC/min.

The furnace was ramped from room temperature to 700oC at 10oC/min in flowing argon. Once at temperature, the furnace was held for 0.5 h before ramping to 1450oC at 10oC/min. A partial vacuum (∼160-120 millitorr) was931pulled and the temperature held for 1 h. Afterward, the furnace was ramped to 1600oC at 10oC/min and held for 1 h. During these high temperature isothermal holds, the mild vacuum was monitored; an increase in pressure indicates volatilization of material or the reaction is occurring. A manual hold, not exceeding an additional 15 min, was performed to allow the vacuum to decrease as much as possible before proceeding. The final ramp from 1600oC to 1900oC at 30oC/min occurred under flowing argon, then dwelled for 2h to sinter, and cooled to 50oC at a rate of 30oC/min.Relative density of each sample was determined using the Archimedes’ method with water as the immersion medium.

All pellets were weighed to obtain their dry mass, D, put in individual beakers of distilled water, and brought to a boil. Then the beakers were transferred to a vacuum chamber, where they were infiltrated for 24 h. The vacuum was checked and re-pulled at 10 h to ensure integrity. After the saturated weight, M, and suspended weight, S, were measured and the bulk density was calculated according to Equation 1 [17]:𝐵 =𝐷𝑀−𝑆(1)Percent relative density is then determined by dividing the bulk density, B, by theoretical density. The value of the theoretical density of a mixed powder is determined by a rule of mixtures calculation.Specimens achieving higher than 98% theoretical density were then mounted to polish the face of the pellet to a 0.25 μm finish using successively finer diamond abrasives.

The specimens were chemically etched in a 1:1 KOH solution at 200oC for 30 s; their microstructures were analyzed using scanning electron microscopy (SEM; field emission S-4700, Hitachi, Schaumburg, Illinois, USA). An average grain size was measured using ImageJ, which is an open source imaging software.2.2 Reaction SinteringThe sequence of reactions in Table 1 was the assumed route of reaction for in situ reaction sintering. Reactions (2) - (5) represent sequential and concurrent reactions, while reaction (1) is the intended overall reaction.

A composition was batched according to the molar ratios of reaction (1) and mixed with WC media in a mill jar for 1 h in MEK. The solvent was removed via rotary evaporation and the powder uniaxially pressed according to the same procedure as the pressureless sintering study. These specimens were reaction sintered using the same rates and isothermal holds as the pressureless sintering study; this is to verify the times and temperatures for reaction. After reaction, the crystalline phases were identified using powder X- ray diffraction (XRD; PANalytical X’pert Phillips, Eindhoven, The Netherlands).2.3 Extrudate BatchingFormulation of 55 vol% 10 mL test batches were done by mixing milled ZrB2 + 4 wt% B4C powders in 50 mL beakers using a flat stainless steel spatula. After achieving a smooth mixture, a binder solution was introduced, of which two types were used for the formulation of extrudate: PVA and Methocel.

After achieving a smooth mixture, a binder solution was introduced, of which two types were used for the formulation of extrudate: PVA and Methocel. A baseline concentration of 1 wt% was used for PVA, while Methocel concentrations were 0.5 and 0.25 wt%. Rheometry was employed to measure the viscoelastic response of each extrudate batch.Due to a lack of applied pressure during fabrication with CODE, in comparison to uniaxial pressing, sintering additives must be included and their effect on the viscoelastic response must also be measured. For deposition of a thick-walled specimen of mesoscale architecture, a shear- thinning rheology is favorable, such that the extrudate behaves like a solid after deposition.2.4 RheometryA digital rheometer (DV-III, Brookfield AMETEK, Inc., Middleboro, Massachusetts, USA) was used for all viscosity measurements, using a small sample adapter and a number 28 spindle.

With these peripherals, the possible viscosity range is 200-5,000,000 cP and the possible shear rate range is 0-56 s-1. The chamber is loaded with 10 mL of extrudate, then the spindle is attached and inserted into the extrudate. Once the rheometer is loaded with a sample, a high shear history is put on the sample by running at 100 RPM for 300 s. Afterward, a viscosity reading is taken every 10 s for 180 s, then the torque is recorded. This is repeated three times for each respective RPM; the RPM is stepped 100, 75, 50, 25, and 5 to provide a range of shear stress.3. Results and DiscussionTwo powders were investigated to confirm the efficacy of pressureless sintering a UHTC: ZrB2 + 4 wt% B4C and ZrB2 + 4 wt% B4C + PVA binder. Compact specimens of each composition were pressed at 70.3 MPa and sintered at 1900oC.

Compact specimens of each composition were pressed at 70.3 MPa and sintered at 1900oC. Pellets pressed without a binder sintered to ∼70% of theoretical density, while specimens containing PVA binder sintered to near theoretical density (≥99%), due to increased green density.933Powders used for dispersion were characterized with surface area and particle size analysis, the results of which are listed in Table 2. For colloidal processing, by definition, some dimension of the powders used must be below one micrometer (< 1 µm). Stability can be achieved using acrylic surfactants to electrosterically disperse particles in suspension, which is greatly affected by particle size distribution, particle geometry, and surface area.Dispersions with Darvan C-N and Dolapix CE 64 all suggest 0.5 wt% Darvan C-N to be most effective at dispersing submicron ZrB2 powder in an alkaline pH≈10 suspension.

Dispersions were checked every 1800 s and most of these rapidly sedimented, leaving a clear or nearly transparent supernatant fluid after 3600 s. The accuracy of this test is in question without zeta potential measurements to verify the dispersion stability; however, it is known that the isoelectric point of ZrB2 is pH ~ 4.5 according to studies done by Leo et.al. [14], and that an alkaline pH > 8.0 will provide maximum stability.Table 2: Measured physical properties of powders.MaterialBoron Carbide (B4C) Boron Nitride (BN) Graphite (C) Zirconia (ZrO2) Zirconium Diboride (ZrB2)Surface Area (m2/g)16.653 9.931 30.172 15.467 1.849Average Particle Size, d50 (µm) 0.82 - - 0.53 2.58Rheometry of the high solids loading test batches with PVA and Methocel binders shows a strong non-Newtonian relationship for batches with 1 wt% PVA and 0.5 wt% Methocel (Figure 1).

The downward trend of decreasing viscosity with increasing shear rates is the hallmark of non- Newtonian response, i.e. shear thinning. One batch of 1 wt% PVA was tested 24 hr after batching, to determine the effect of sedimentation and current extrudate shelf-life. At the lowest shear rates, the drag on the spring-spindle assembly reached the maximum torque; similar behavior is seen with low concentrations of Methocel. This response is due to high shear stress from friction between the sediment, in the case of the PVA solution, and due to low concentrations of Methocel, which would otherwise act as a plastic barrier between particles. Higher concentrations of binder exceeded the maximum viscosity of the rheometer, while lower concentrations did not provide consistent data, trial to trial, across the entire shear stress range.

Higher concentrations of binder exceeded the maximum viscosity of the rheometer, while lower concentrations did not provide consistent data, trial to trial, across the entire shear stress range. An alternative method to measure the viscoelastic response of high solids loading suspension would be favorable to fully characterize viscoelastic response with a more representative range of additive concentrations.Microstructural analysis was done to measure the average grain size and verify the sintering schedule of pressurelessly sintered ZrB2 + 4 wt% B4C. As shown in Figure 2, there is no intergranular or intragranular porosity, which verifies the sintering time and temperature to achieve high relative density. In the high magnification image (a), it is possible to see some voids, but this is due to polishing and the caustic nature of the KOH etchant, not a product of sintering.

In the high magnification image (a), it is possible to see some voids, but this is due to polishing and the caustic nature of the KOH etchant, not a product of sintering. A grain934size analysis was done using ImageJ to measure the average grain area and estimate an equivalent circular diameter. Using this method, the approximate grain diameter of the specimens sintered at 1900oC is 10.56 µm, based on a count of 797 grains. This included only the grey phase, in Figure 2, which is ZrB2, while the black phase is B4C.(a)(b)Figure 1: Comparison of PVA and Methocel binders in high solids loading batches.Figure 2: Microstructure of pressurelessly sintered ZrB2; (a) high magnification showing no inter- or intragranular porosity; (b) low magnification, showing grain development and pinning.935A reaction to yield ZrB2 is often the decomposition of a boron species (borothermal) or carbon species (carbothermal) with a zirconium source.

The oxide-carbide-nitride system proposed is a hybrid boro/carbothermal reduction. This multi-step reaction (Table 1) was verified by thermochemical software, however the path may be effected or hindered by particle size, the isothermal holds during sintering, and partial vacuum steps during sintering. For this reason, powder XRD analysis of crushed samples was used to identify the product, revealing that the predominant crystallographic phase of the powder sample was zirconium diboride (Figure 3a). Residual phases from incomplete reaction would present themselves as intermediate intensity peaks, such that peaks in Figure 3b rows 3-5 would be present in rows 1. However, this is not the case, as the peaks of the scanned sample, Figure 3a and 3b row 1, correspond to the chemical tag of zirconium diboride (Figure 3B row 2).

However, this is not the case, as the peaks of the scanned sample, Figure 3a and 3b row 1, correspond to the chemical tag of zirconium diboride (Figure 3B row 2). None of the major peaks of the reactants, Figure 3b rows 3-5, present themselves in Figure 3a or 3b row 1, confirming the formation of single phase crystalline zirconium diboride.Figure 3: XRD Micrograph of predominant phase of reaction sintered ZrB2.4. Conclusion and Future WorkOur goals stated within the scope of this study, 1) identify viable reaction materials, and 2) determine the ratio of additives to promote pressureless sintering to achieve high relative density, have been achieved. A relative density ≥ 99% was achieved by pressurelessly sintering milled ZrB2 + 4 wt% B4C at 1900oC, and a successful oxide-carbide-nitride system was established to yield ZrB2 with an average grain size of 10.56 µm.

Two aqueous binders were presented as options for developing an extrudate for reaction sintering; both PVA and Methocel show a high degree of shear thinning at low weight percentage, which is ideal for CODE deposition.936Further study is necessary to fully characterize the extrudate via rheometry; the storage and loss moduli are of particular interest for non-Newtonian extrudate. Having confirmed a chemistry for reaction sintering, further development is necessary to achieve high relative density. Once achieved, green ceramic parts will be fabricated using CODE for mechanical testing and microstructural analysis.AcknowledgmentsThe authors gratefully acknowledge the financial support by the National Energy Technology Laboratory of the Department of Energy under the contract #DE-FE0012272, and by the Intelligent Systems Center at the Missouri University of Science and Technology.References[1] X. Zhang, J. Han, and Z.

DESIGNER'S GUIDE FOR DIMENSIONING AND TOLERANCING SLS PARTSCarolyn Conner Seepersad, Tyler Govett, Kevin Kim, Michael Lundin, Daniel Pinero Department of Mechanical Engineering University of Texas at AustinABSTRACTBecause additive manufacturing (AM) is a relatively novel industry, with the first commercial machines introduced in the late 1980s, many designers are unaware of the capabilities of AM technologies. Many engineers also find it difficult to utilize AM because of a lack of “Design for AM” knowledge in the public domain. Reliable information on material properties, dimensions and tolerances, and other process-related specifications is often scattered throughout the literature, if it is publicly available at all. The objective of the research reported in this paper is to begin to create a designer's guide for dimensioning and tolerancing parts that are additively manufacturing using selective laser sintering (SLS) technology.

The guide is based on a series of experiments designed to determine the limiting feature sizes for various types of features fabricated in commercially available SLS machines. The features include slits, holes, letters, mating gears, and shafts built in a preassembled state. The impact of part thickness, orientation, clearance, and dimensions on the resolvability of features is examined. Results are reported in a series of matrices that relate realizable feature sizes to other important variables such as part thickness.Keywords: Additive Manufacturing, selective laser sintering, tolerancing, dimensioning, design for additive manufacturing1. OVERVIEW OF A DESIGNER'S GUIDE FOR DIMENSIONING AND TOLERANCING SLS PARTSAM is attractive to engineering designers because it offers unique capabilities that are not found in conventional subtractive or formative manufacturing processes [1].

Using selective laser sintering (SLS), designers have the opportunity to fabricate almost any shape or topology, including complex internal structures such as cellular or lattice structures. These capabilities can be used to consolidate components for lightweighting and ease of assembly, to tailor customized structures for form-fitting and multifunctional applications, to fabricate moving joints and mechanical assemblies in situ, and to create unique, one-of-a-kind products. Without the need for dedicated tooling, these products can be fabricated economically in lot sizes as small as one.It can be difficult, however, for experienced or novice engineers to design for AM. The expansive capabilities of AM the design-for-manufacturing (DFM) constraints of conventional manufacturing processes, thereby expanding the design freedom of the designer, but they also leave the designer with new sets of process-specific design rules that are often poorly understood and quantified.

Accuracy and resolution of different features can vary from AM process to process and also with part orientation, thickness, and other characteristics of the part and the build.lift921Without detailed, quantitative knowledge of the accuracy and resolution capabilities of a specific AM process, defects are common. For example, Figure 1 displays a selection of products fabricated in SLS by undergraduate and graduate students in an AM course at UT Austin. Defects include unreadable text, thin walls that lack integrity, and moving parts that either fuse together or fail to mate properly.Figure 1. Examples of defects in SLS parts.

Defects include unreadable text, thin walls that lack integrity, and moving parts that either fuse together or fail to mate properly.Figure 1. Examples of defects in SLS parts. L to R: unresolved walls, unreadable text, slipping gearsAlthough expert AM part designers may be able to avoid these defects by leveraging their extensive experience fabricating similar features, it is important for less experienced AM part designers to have access to publicly available information on the accuracy and resolution of AM processes with respect to common sets of features. It is also important for this information to be much more detailed and feature-specific than the accuracy and minimum feature size estimates that are typically available from manufacturers. Towards this goal, several authors have designed and fabricated benchmark parts and published comparative studies of prominent processes such as SLS, stereolithography (SLA), and fused deposition modeling (FDM) (e.g., [2- 4]).

These studies typically investigate the accuracy and repeatability of a variety of features, including cubes, cylinders, slots, holes, and overhanging beams, and also investigate warpage, curl, and surface finish under various conditions.These benchmarking studies leave several unresolved questions that can be important for designers of mechanical parts for SLS.

Those questions include: How much should mating gears be separated to prevent both fusion of the mating teeth and slipping between the moving teeth? How much clearance should be provided between a rotating shaft and a surrounding bore if the parts are fabricated in an assembled state? How does the answer differ with orientation and thickness of the surrounding part? How does the resolvability of holes and slots and thin walls vary with orientation and the thickness of the surrounding part? What types and sizes of fonts are appropriate for SLS and how does the answer change with orientation of the surface and the raised or indented nature of the font?In this paper, a set of benchmark parts are designed, fabricated, and measured to answer these questions for the 3D Systems Vanguard series of SLS machines. A subset of benchmark parts are introduced in Section 2, along with a sampling of results in Section 3.

A subset of benchmark parts are introduced in Section 2, along with a sampling of results in Section 3. A full description of the benchmark parts and results is available in a full-length report available on the first author's website [5].9222. BENCHMARK PARTSBenchmark parts were designed to investigate feature resolution, font resolution, and clearance between moving mechanical parts. Most features were fabricated on parts of varying thickness and with different build orientations to investigate the effects of those factors on resolution and accuracy. Separate benchmark parts were designed for each type of feature to be investigated.

Separate benchmark parts were designed for each type of feature to be investigated. Many of the designs (along with the graphical approach for presenting the resulting measurements) were inspired by the research of Dominik Sippel, as described on the Shapeways website [6]; however, his work investigated only EOS laser sintering machines, rather than 3D Systems Vanguard machines, and did not investigate the moving mechanical parts described here. A selection of the benchmark parts is described in this section, which is organized by benchmarking category.2.1 Feature ResolutionFeature resolution experiments included the investigation of circular hole resolution versus part thickness, circular hole resolution versus proximity to the edge of a part, thin wall resolution, small diameter pin resolution, and square hole resolution versus part thickness.

All feature resolution experiments were conducted in two orthogonal orientations in which the primary plane of the part was either parallel or orthogonal to the build plane.Two example types of benchmark parts are pictured in Figures 2 and 3. In each figure, the z- axis is aligned with the height of the build chamber, and the paired pictures illustrate the two orientations in which the parts were fabricated. Figure 2 illustrates the benchmark part for investigating circular hole resolution versus part thickness. The circular holes ranged in diameter from 0.125 mm to 4.0 mm, and the part thickness ranged from 0.94 mm to 12.7 mm. As shown, identical parts were built in horizontal and vertical orientations in which the central axis of the holes was aligned with the height and width of the build chamber, respectively. Figure 3 illustrates the benchmark part for investigating thin wall resolution versus part thickness. The thin walls ranged in thickness from 0.2 mm to 3 mm.

Figure 3 illustrates the benchmark part for investigating thin wall resolution versus part thickness. The thin walls ranged in thickness from 0.2 mm to 3 mm. The part was build in two orientations, as shown, with the thin walls either coplanar with the build plane (horizontal orientation) or orthogonal to the build plane (vertical orientation). Two duplicate parts were fabricated in each orientation for each part.Horizontal OrientationVertical OrientationFigure 2. Benchmark Part for Circular Hole Resolution Versus Part Thickness923Horizontal OrientationVertical OrientationFigure 3. Benchmark Part for Thin Wall Resolution Versus Orientation2.2 Font ResolutionFont resolution experiments included fabricating parts with different fonts in different orientations and with different font sizes and heights (if raised) or depths (if recessed). As shown in Figure 4, grids of letters were fabricated with different font sizes and heights/depths.

As shown in Figure 4, grids of letters were fabricated with different font sizes and heights/depths. On each surface, both raised and recessed letters were fabricated. Letters were fabricated on both sides of the plate, as well, to investigate upward and downward facing font resolution for the horizontal orientation shown in Figure 5. The vertical orientation in Figure 5 is used to investigate the resolution of side-facing font.Figure 4. Benchmark Parts for Two Different Fonts (Sans Serif on the Left and Serif on the Right). Font sizes vary from Left to Right. Font Heights/Depths Vary from Top to Bottom. Raised Letters are Pictured on the Upper Left; Recessed Letters on the Lower Right.924Horizontal OrientationVertical OrientationFigure 5. Build Orientations of Benchmark Parts for Font.2.3 Clearances for Moving Mechanical PartsClearance experiments included shafts with varying bore diameters and gear assemblies with varying clearances.

Figure 6 illustrates the benchmark part for investigating shaft clearance. Part thickness decreases from 25.4 mm to 1.27 mm along the y-axis, and bore diameter decreases incrementally from 13.6 mm to 10 mm along the x-axis. Shafts are identically dimensioned with 4 mm lengths and 10 mm diameters. The part was fabricated in only one orientation, as shown in Figure 6, with the central axis of each shaft aligned with the height of the build chamber (the z-axis).Figure 6. Benchmark Part for Shaft Clearance.Gear clearances were investigated with the benchmark part in Figure 7. As shown in the figure, six pairs of gears were fabricated on each test part. Shaft/bore clearances of either 1 or 1.5 mm were tested, together with gear tooth clearances of 0.5, 1, or 1.5 mm. Three different versions of the test part were created with 15, 20, and 25 gear teeth per gear.

Three different versions of the test part were created with 15, 20, and 25 gear teeth per gear. The part was always aligned as shown in the figure, with the z-axis aligned with the height of the build chamber, and two identical copies of each part were fabricated for repeatability.925Figure 7. Benchmark Part for Investigating Gear Clearances3. REPRESENTATIVE RESULTSAll of the benchmark parts were fabricated on a 3D Systems Vanguard SLS machine by Harvest Technologies. Harvest Technologies used production machine settings and Nylon 12 powder. Metrology was conducted on each fabricated part. Pass/fail criteria for each feature were compiled to determine whether the feature was adequately resolved. Tolerances were also quantified by measuring the parts with calipers for larger features or an optical microscope for smaller features. Each measurement was repeated by four members of the team, and results were averaged to improve accuracy and reduce bias from individual measurements.

Each measurement was repeated by four members of the team, and results were averaged to improve accuracy and reduce bias from individual measurements. A representative sampling of the results is provided in this section, with full results available in the original report [5].3.1 Feature ResolutionThe pass/fail criteria for circular holes are described in Figure 8. Results for a vertically oriented part and a horizontally oriented part are illustrated in Figure 9, with the orientations corresponding to Figure 2. As shown in Figure 9, hole resolution is better for parts oriented in the vertical orientation; as shown in Figure 2, vertically oriented parts include holes with their central axes aligned with the build plane. For those parts, the thickness of the build layer has a stronger influence on the resolvable hole size than the beam width.

For those parts, the thickness of the build layer has a stronger influence on the resolvable hole size than the beam width. Since the layer thickness for these builds was approximately 0.1 mm, while the beam width was approximately 0.28 mm, it is not surprising that smaller holes are resolved for the vertically oriented parts. Part thickness also has an influence on hole resolution, with smaller holes resolvable in thinner parts. Thicker parts require more laser scanning in the region of the hole, likely resulting in small amounts of oversintering and coarser hole resolution. In addition to these charts of hole resolution, hole diameters were measured repeatedly, and resulting statistics are documented in the full report [5].926Figure 8. Pass/Fail Criteria for Circular Holes.Figure 9. Circular Hole Resolution Versus Plate Thicknesses for Vertical Part Orientation (Left) and Horizontal Part Orientation (Right). See Figure 2 for Orientation Description.

Circular Hole Resolution Versus Plate Thicknesses for Vertical Part Orientation (Left) and Horizontal Part Orientation (Right). See Figure 2 for Orientation Description. Color coding matches Figure 8.The pass/fail criteria for thin walls are described in Figure 10. Results for a vertically oriented part and a horizontally oriented part are illustrated in Figure 11, with the orientations corresponding to Figure 3. As shown in Figure 11, thin wall resolution is better for parts oriented in the vertical orientation, in which the walls are coplanar with the build plane. For parts in the preferred orientation, it is possible to build walls as thin as 0.2 mm (less than twice the layer thickness), although those walls are very thin—so thin that they buckle under the residual stresses associated with the cool-down of the build. In either orientation, it is possible to resolve walls as thin as 0.8 mm.927Figure 10. Pass/Fail Criteria for Thin Walls.Figure 11.

In either orientation, it is possible to resolve walls as thin as 0.8 mm.927Figure 10. Pass/Fail Criteria for Thin Walls.Figure 11. Thin Wall Resolution Versus Wall Thickness for Vertical Part Orientation (Top) and Horizontal PartOrientation (Bottom). See Figure 3 for Orientation Description. Color Coding Matches Figure 10.3.2 Font ResolutionThe pass/fail criteria for fonts are described in Figure 12. Results for a horizontally oriented part are illustrated in Figure 13, with the orientation corresponding to Figure 5. As shown in Figure 13, font resolution is much better for upward facing surfaces than for downward facing surfaces. The downward-facing letters may suffer from oversintering associated with building thick parts around and on top of the letters.Figure 12. Pass/Fail Criteria for Fonts.928Figure 13. Font Resolution for Serif Font, Recessed Letters, and Horizontal Part Orientation.

Pass/Fail Criteria for Fonts.928Figure 13. Font Resolution for Serif Font, Recessed Letters, and Horizontal Part Orientation. Font Size is Charted on the Horizontal Axis Versus the Depth of the Recessed Letter on the Vertical Axis. Downward Facing Letters are Depicted on the Left; Upward Facing on the Right. Color Coding Matches Figure 12.3.3 Clearances for Moving Mechanical PartsThe pass/fail criteria for gears are described in Figure 14. The acceptability of the mesh between the gears is illustrated in Figure 15. In Figure 15, failures (color-coded in red) indicate that the gear teeth are too far apart to maintain reliable contact; none of the gears fused together in this experiment. Based on these results, it is reasonable to recommend a shaft clearance of 1 mm (but no larger) and a gear tooth separation of 0.5 to 1.0 mm.Figure 14. Pass/Fail Criteria for Gears.929Figure 15.

Pass/Fail Criteria for Gears.929Figure 15. Acceptability of Meshing Gears as a Function of Separation Distance Between Gear Teeth, Clearance Between Gear Inner Diameter and Shaft, and Number of Gear Teeth. Gears are Oriented as Shown in Figure 7.4. CLOSUREThis paper offers a snapshot of the of the results of a series of experiments designed to determine the limiting resolution of a selection of features fabricated in commercial SLS machines. The matrices provided in this paper relate feature resolution to other important parameters such as part thickness, clearances, and build orientations.The complete designer's guide offers additional matrices for features including shafts preassembled within a bore, holes in proximity to the edge of a part, slits, and many different permutations of the results reported here, such as additional build orientations.

The complete designer's guide also includes quantitative analysis of the accuracy of specific features based on repeated measurements of that feature in multiple parts and multiple build orientations.The results are certainly limited in several ways. First and foremost, the results are specific to the 3D Systems Vanguard HiQ+HiS machine and Nylon 12 powder. Other machines and/or materials could yield different results. Secondly, many features remain unexplored, such as springs, fasteners, and knobs, along with geometric characteristics such as warping and concentricity.ACKNOWLEDGEMENTSThe authors gratefully acknowledge Harvest Technologies for sponsoring the project and fabricating all of the test parts. Thanks also to the Senior Design Projects program in UT Austin's Mechanical Engineering Department for making the project possible.930REFERENCES1. Bourell, D. L., M. Leu and D. W.

S. Wong, J. Y. H. Fuh and H. T. Loh, 2004, “Benchmarking for Comparative Evaluation of RP Systems and Processes,” Rapid Prototyping Journal, Vol. 10, No. 2, pp. 123-135.5. Govett, T., K. Kim, M. Lundin and D. Pinero, 2012, "Design Rules for Selective Laser Sintering," Senior Design Project Report, Mechanical Engineering Department, The University of Texas at Austin. Available at www.me.utexas.edu/~ppmdlab [Accessed August 17, 2012].6. “Design Rules and Detail Resolution for SLS 3D Printing,”http://www.shapeways.com/tutorials/design_rules_for_3d_printing [Accessed August 17, 2012].931.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDesigning a Power Tool to Show the Potentials of Additive Manufacturing - Effects of Additive Manufacturing on the Product Development ProcessB. Leutenecker-Twelsiek*, C. Klahn †, and M. Meboldt*ETH Zurich, Leonhardstrasse 21, 8092 Zurich, Switzerland †inspire AG, Leonhardstrasse 21, 8092 Zurich, SwitzerlandAbstractToday Additive Manufacturing (AM) is mainly used for rapid prototyping and specialized parts for industrial and end-user applications. To communicate the design potentials of AM to a general audience it is useful to demonstrate it with a visionary product, similar to the concept cars of the automotive industry. We have redesigned a power tool to show the benefits of AM on an end-user product.

We have redesigned a power tool to show the benefits of AM on an end-user product. This paper describes the Concept Tool, its different sub-systems and highlights the range of AM’s applications. Furthermore we present our observations and findings during the product development process. Based on these findings a design supporting system is suggested to improve the development process for additive manufactured products.IntroductionAdditive Manufacturing (AM) [1] processes are technologically mature for industrial production due to a sufficient process stability, productivity and resulting part quality. A rising competition between service providers [2] makes Additive Manufacturing economically feasible for a growing number of industrial and end-user applications [3]. Nevertheless, companies still face a number of challenges when implementing AM in their production of end-user parts. First the companies have to identify suitable parts for AM [4].

Nevertheless, companies still face a number of challenges when implementing AM in their production of end-user parts. First the companies have to identify suitable parts for AM [4]. The second big challenge is to (re)-design the identified part, which was designed for conventional manufacturing, to allow an economical AM production.AM is a new production process and allows new possibilities in design, but also restricts the design in different way than conventional manufacturing. The key to unlock the design potential of AM is to overcome the traditions and fixed mind-sets of experienced designers and of course the lack of knowledge on the new technology [5,6].Design showcases are a useful aid to understand the potentials of AM for new products. Reflecting the showcases helps a designer to identify suitable parts with similar potentials on the product portfolio of their own company.

Reflecting the showcases helps a designer to identify suitable parts with similar potentials on the product portfolio of their own company. There are various publications [3,6-8] presenting collections of individual parts to show different potentials of AM. To enable audience to understand the benefits of AM based on individual parts one has to explain its function, the surrounding context and application. One can save these tedious explanations by implementing the AM benefits into an easy to understand product. In the automotive industry they implement future technologies in only one showcase, commonly referred to as concept cars. Based on this procedure, the idea was born not to develop a new part but rather to develop a completely functional system to illustrate the different potentials of AM.

Based on this procedure, the idea was born not to develop a new part but rather to develop a completely functional system to illustrate the different potentials of AM. Furthermore, observation of the development process of a complete system allows tenables identification of challenges and implications of AM on the overall product development process.1900To avoid any bias of this observation from existing AM knowledge a team of AM novices was selected. The development project of the Concept Tool was started in the framework of a Focus Project at ETH Zurich. An interdisciplinary team with six mechanical engineering, two electrical engineering and two industrial design students had to develop a working system within nine months.

An interdisciplinary team with six mechanical engineering, two electrical engineering and two industrial design students had to develop a working system within nine months. The team selected a specific of handheld power tools and developed an additive manufactured tool to illustrate different potentials of AM.To enable the team to select a suitable power tool they were introduced in the potential of AM and the four selection criteria, namely integrated designs, individualization, lightweight design, and efficient design [9]. With the selection criteria the team analyzed different power tools and decided on battery operated hammer drill. A list of requirements was specified for the hammer drill with an emphasis on demonstrating AM potentials. The mechanics and the working principles of a hammer drill had to be entirely understood to select the optimal functions or components for improvements and substitutions.

The mechanics and the working principles of a hammer drill had to be entirely understood to select the optimal functions or components for improvements and substitutions. Therefore, a structural functionality flowchart and a thorough system definition were made. In a next step a potential analysis was performed to determine which functions and sub-assemblies have the most potential to illustrate the benefits and capabilities of AM technologies. Six different models of drilling machines were disassembled and documented for this analysis. Some of the tools had additional gadgets, for example a dust extraction system. Each sub-assembly was assessed with respect for the four selection criteria.Figure 1 plots the potential of each sub-assembly over the expected risk and effort of designing AM-solution challenge. The ideal part would have big potential and it would be a small challenge to realize.

The ideal part would have big potential and it would be a small challenge to realize. The best suited functions and sub-assemblies for the purpose of the Concept Tool were a gear box, a dust extraction system, the housing, the unit to generate the percussion energy and some electronic features like a horizontal positioning aid or borehole depth setting or.

Other components like gear or electric motor where put out of scope for a re-design, because of their unfavorable ratio between challenge and potential benefits.Figure 1: Evaluation of potential analysis1901Based on the outcome of the function and potential analysis the team developed the following Concept Tool in an iterative process.Concept ToolThe final design of the Concept Tool (see figure 2) offers all functions of a conventional hammer drill and some more like integrated dust extraction, digital controlling with display and novel internal device structure.Figure 2: Final design of the Concept ToolThe Concept Tool is able to drill into wood and metal, hammer drill into concrete and stones or tighten screws. In the hammer drilling or drilling mode the desired bore hole depth can be set using a selection wheel placed in the front (see .figure 3). If the mode is set to screwing the torque can be adjusted using the same wheel.

If the mode is set to screwing the torque can be adjusted using the same wheel. The machine stops automatically if the depth or torque limit is reached. An integrated display indicates the selected depth or torque. Furthermore, the direction of rotation and the mode are indicated using illuminated symbols. A main feature of the Concept Tool is the dust extraction system. The drill dust gets vacuumed in at the base of the tool and is transported to an integrated filtering system. The dust accumulates at the bottom of the handle, inside a dust collector. The collector can be removed, emptied and reattached to the tool. Before the cleaned air leaves the tool it is used to cool the electric motor, as well as to blow the drill dust into the extraction base.

Before the cleaned air leaves the tool it is used to cool the electric motor, as well as to blow the drill dust into the extraction base. If the dust extraction system is not needed the base can be locked at the tool by pulling it all the way to the chuck and moving the slider on top of the base to the left.Figure 3: Selection wheel and rotary switch (left), display (middle), dust extraction base with aperture (right)1902The Concept Tool is powered by a Lithium Polymer (LiPo) battery. The battery is exchangeable and is locked to the tool using a newly developed magnetic locking mechanism. It snaps automatically to a locked position and mechanically fixes the battery to the tool.

The battery is exchangeable and is locked to the tool using a newly developed magnetic locking mechanism. It snaps automatically to a locked position and mechanically fixes the battery to the tool. An SDS Quick mechanism from Bosch allows a fast and easy way of exchanging the drills.Figure 4: Magnetic lock slider integrated in the housing of the exchangeable battery (left), pressure sensor covered with an flexible grid material (middle), horizontal positioning aid (right)The Concept Tool is controlled in a unique way only with the index finger. All the main functions can be controlled by applying force on pressure sensors placed on both sides of the tool. It is therefore designed for left- and right-handed people. With a double tap on the pressure sensor the direction of rotation can be changed. Slight pressure with the index finger triggers a laser line projected to the wall.

With a double tap on the pressure sensor the direction of rotation can be changed. Slight pressure with the index finger triggers a laser line projected to the wall. An electric version of a spirit level indicates with three Light-Emitting Diodes (LEDs) if the tool is held horizontal or not (see figure 4). This horizontal positioning aid allows drilling multiple holes at the same height. By firmly pressing the index finger on the pressure sensor starts the drilling process. The speed of rotation is controlled by the amount of force applied by the index finger. A sudden increase in pressure triggers the safe mode of the tool and drilling is impeded. This prevents unwanted initiations of the drilling process, for example when the tool falls to the ground.With the exception of the bearings, shafts, gears and screws, all designed parts are produced with AM. The motor cooling unit and the bearing structure are produce by Selective Laser Melting (SLM) out of aluminum.

The motor cooling unit and the bearing structure are produce by Selective Laser Melting (SLM) out of aluminum. The other parts of the Concept Tool are produced by Selective Laser Sintering (SLS). The centrifugal fan are made of Duraform HST, the grips are made of Duraform flex and all other designed parts are made of PA12.Subsystem bearing structureThe Concept Tool’s innovative and AM specific solutions demonstrate the possibilities of AM. Three of the solutions are highlighted and present here. The gear and bearing in hammer drills are normally stored in an exoskeleton more precisely in the housing of the power tool. To show the potential of lightweight design in AM a bearing structure was designed based on the idea of creating a gear subassembly that absorbs all gear forces within the assembly. The goal was to optimize the structure with regard to predefined load cases and use as little material as possible.

The goal was to optimize the structure with regard to predefined load cases and use as little material as possible. Such complex, weight optimized geometries are usually not manufacturable with conventional processes. In contrary to conventional manufacturing the complexity of the geometry does not affect the costs1903of additive manufactured parts. The bearing structure is therefore a very suitable part to show what AM is capable of in terms of lightweight design of load bearing parts.The final design of the bearing structure is a scaffold with a half shell and four additional parts to fix the gear (see figure 5). This ensures an easy and fast way to assemble the gears and bearings, which are taken from a conventional power tool. The build direction of the structure was taken into account so it was possible to place the bearing seats in such a way that little to no postprocessing is necessary.

The build direction of the structure was taken into account so it was possible to place the bearing seats in such a way that little to no postprocessing is necessary. Furthermore by applying the principle of early determination of the part orientation [10] it is possible to manufacture the bearing structure with a minimum of support material. The main part of the gear subassembly is built in the orientation shown in figure 5. The large, flat additive manufactured sheet at the bottom is necessary and acts as an interface between the structure and the cooling system. The other four additional parts of the subassembly are oriented differently to also allow manufacturing without any additional support material. At the end of the build process the few necessary supports can be removed easily. Besides heat treatment and sand blasting of all parts, only four threads are cut during postprocessing.

At the end of the build process the few necessary supports can be removed easily. Besides heat treatment and sand blasting of all parts, only four threads are cut during postprocessing. The minimal weight achieved with an initial conceptual design based on the flux of the forces around the gears. Based on this ideal design a force optimized part was developed in several iterations of simulations and design changes. When hammer drilling with at a force of 100 N and a torque of 0.8 Nm the maximum stress in the final part is approximately 72 N/mm2 and the maximum displacement 0.18 mm with a total weight of 60g.Figure 5: Design of the gears and bearing structure with the four additional parts (left), assembled part with gears and bearings (right)Subsystem dust extraction and motor coolingThe goal of an integrated dust extraction system is to ensure a clean drilling process. The Concept Tool should be able to drill a hole with no visible dust left on the wall or below on the floor.

The Concept Tool should be able to drill a hole with no visible dust left on the wall or below on the floor. The dust extraction system is integrated into the drilling machine such that it does not need any additional equipment like a vacuum cleaner. The system is able to filter the dust out of the air1904and the user of the tool can easily empty the filter system. To realize this feature a complex ventilation and extraction system was developed with a cyclonic separator, centrifugal fan and directly at the drilling hole an air outlet and the extractor inlet. The air flow is also used to cool the electric motor (see figure 6).Figure 6: Overview of ventilation and dust extraction system (above), cyclonic separator (left), centrifugal fan with motor cooling (middle) and dust extraction base (right).The heart of the dust extraction system is the centrifugal fan. It provides the necessary airflow and static pressure.

It provides the necessary airflow and static pressure. The negative pressure at the inlet is used for dust extraction and transports the dust directly to the cyclonic separator. In the cyclonic separator the heavy dust particle are separated and fall down into the dust collector at the lower end of the handle. The collector is attached with a thread and easy to remove. When the cyclonic separator is turned upside down, a stopper in the dust collector prevents the dust from falling back into the separator. To protect the centrifugal fan and the motor cooling from remaining dust particles, a fine particle filter was added directly above the cyclonic separator. In this configuration, particles in the fine particle filter can fall down into the dust collector when the dust extraction system is turned off.The cleaned air flows into an additive manufactured cooling structure for the motor.

The cooling structure is optimized for an optimal flow and heat exchange and is not manufacturable in a conventional way. It is built in aluminum without any support material. The optimized airflow reduces the motor temperature of ten Kelvin. After the cooling system the majority of the air leaves the tool through an outlet. A small portion is redirected to the front into the base. This airflow is used to clean the drill and the drilling hole. The outlet is placed opposite of the inlet of the dust extraction channel. The channels to transport the clean and dusty air to and from the drilled hole are integrated in the beams holding the dust extraction base. The cooling system and the whole1905flow optimized dust extraction system are a good illustration of how AM can increase the performance with an optimal design.The dust extraction system is further enhanced by another interesting feature in the dust extractor base.

To ensure the vacuum cleaning of the drilling hole, it is important that the dust extraction system draws air and dust from the hole and not the ambient air. The backside of the base has to be closed around the rotating drill. To enable different drill diameters the closure on the backside has to be adjustable. Therefore the adjustable aperture in figure 7 was designed. The base contains four movable parts, three blades and a control ring, but was manufactured by selective laser sintering directly inside the base in one piece (see figure 7).Figure 7: Aperture: manufactured in closed position (left), sectional view (middle/right)Thes three blades are guided in the base at one end and in a movable ring at the other end. The ring can be moved with the slider to open and close the blades of the aperture.

The ring can be moved with the slider to open and close the blades of the aperture. The integrated aperture with its printed movable part demonstrates functional integration in additive manufactured parts.There are numerous other good ideas illustrating the different potentials of AM in a highly integrated holistic system. The general function of the hammer drill does not need an explanation for a general audience to be understood. It is a perfect visualization of AM’s potentials and allows direct discussion of them in several subsystems and solution details.Reflection of the design processThe whole nine month of the product development process were highly influenced by the different capabilities of additive manufacturing in comparison to conventional processes.

The short lead time of AM parts allowed more and faster design iterations.The main technical functions were treated as individual subprojects following its own individual plan and pace, only bound to the overall project deadlines, like design freezes. Each subproject was realized by a sub-team consisting of up to three members. The engineering boundaries for each sub-project were the interfaces and predefined installation spaces.This adaptive project organization was also needed because the various AM processes bring different production speeds and costs. While metal parts are relatively expensive and need to be1906scheduled with a production period of up to two weeks, some rapid prototyping plastic parts are available within a couple of hours and relatively cheap. Therefore, the realization process of each subproject was bound to the corresponding manufacturing method.

Therefore, the realization process of each subproject was bound to the corresponding manufacturing method. Other subprojects such as the development of the electronic system followed their individual iteration steps. Nevertheless, all sub-projects, of mechanical or electronic nature, were similarly organized. The progress of the functions and their feasibility were evaluated in regular meetings. If needed, a non-realizable function was dismissed and replaced. For example no feasible solution was found in adequate time for a hammering mechanism. Therefore the components of an existing hammer drill were reused (see figure 8). This chosen project structure allowed to continuously test, evaluate, extend and improve single components and consequently the complete product.Figure 8: Development processAdditive manufacturing enabled the large number of iterations, but AM was also a cause for iteration. The team members were novices in the area of additive manufacturing.

The team members were novices in the area of additive manufacturing. They were trained with the basic knowledge on AM and had access to common and available design guidelines. The team designed the first realization of the tool based on this knowledge. During the assembly of the parts and the following tests, valuable feedback was derived from the additive manufactured parts. The feedback was used to improve the design, especially to reduce postprocessing and to fix not working functions.During course of the project it turn out there is an additional need for supporting different aspects of design process with an AM design guideline, particularly for AM novices, was determined. The inadequate designs of the first realization were primarily traced back to insufficient information given by design rules and design guidelines for specific design situations.

The inadequate designs of the first realization were primarily traced back to insufficient information given by design rules and design guidelines for specific design situations. The tow following situations are exemplary for these problems.The first exemplary situation is the need to remove remaining powder from cavities is already know and described. Especially in Selective Laser Sintering the removal can be difficult. In several1907publications of design rules numerical values are given on the minimal gap between two walls. The given information is correct and important, but it corresponds to simple geometries and does not have to be valid for more complex ones. There is no additional information for complex geometry and how to design a part for an effective and efficient powder removal.A second exemplary situation was the lack of guidelines for the overall design process. The team started with an idea and designed the part in CAD.

The team started with an idea and designed the part in CAD. After feedback for the AM service provider and postprocessing they had to redesign the part. In particular on metal parts produced by Selective Laser Melting (SLM) one can save support material and postprocessing effort and in the end to save cost, by an improved design process. Based on this learning we started to develop a guideline for the design process. For example, the part orientation should be determined before designing the part in CAD. This enables the incorporation of the orientation in the detailed design and reduces the need for support material, improves surface quality and reduces heat deformation. Due to the early determination the effort for postprocessing is reduced and design iterations can be avoided [10].ConclusionFor a better understanding and visualization of AM’s possibilities a working system - the Concept Tool - is presented.

The presented Concept Tool has a various features, which demonstrate the potential and design freedom of AM. Especially, features like the flow channels for dust extraction and aperture are only producible with AM. The findings of our team observations showed a lack in supporting systems for designer and emphasized the need for a supporting design guideline for the overall development process. The goal is to support designers in the design process of systems containing AM components and to unlock the full potential of AM.AcknowledgementsThe authors would like to thank the Team ADDit for the development of a unique and inspirational product – the Concept Tool. Without their help, and their input this paper would not have been possible. Oliver Kirchhoff, Manuel Heckhorn Ghilardi, Jan Zimmermann, Helen Meyer, Michael Liem, Mickey Fürer, Thilo Weber, Rolf Scheuner, Pascal Trachsler, Alessio Maag.References[1] ISO/ASTM 52900, 2015-12-00.

signing and slicing heterogeneous components for Rapid PrototypingAshok V. Kumar and Jongho Lee Department of Mechanical Engineering, University of Florida, Gainesville, FL 32611 USAAbstractMany rapid prototyping techniques have the potential for fabricating components whose composition is non-uniform and varies in a desired fashion. A shape and composition modeling technique was developed to enable the representation and design of such heterogeneous components. Techniques for interactively and automatically designing such components are presented. Automatic design is made possible using optimization techniques where the optimal composition distribution is computed based on specified design objective and constraints. Software was also developed to slice 3D heterogeneous solids to generate cross-sectional images as well as composition distribution for each cross-section.

Software was also developed to slice 3D heterogeneous solids to generate cross-sectional images as well as composition distribution for each cross-section. Slicing and generation of cross-sectional data are essential to enable rapid prototyping of these components.1. IntroductionHeterogeneous components are parts whose material composition varies from point to point within the part. Such components are difficult to fabricate using traditional methods. Rapid prototyping systems build parts by adding material layer-by-layer and are in principle capable of varying the composition within manufactured parts.

Rapid prototyping systems build parts by adding material layer-by-layer and are in principle capable of varying the composition within manufactured parts. Progress towards composition control of heterogeneous components have been reported recently using SFF processes such as Selective Laser Sintering (SLS) (Jepson et al, 1997), 3D printing (Yoo et al, 1998), Shape Deposition Manufacturing (SDM) (Fessler et al, 1997), Direct Light Fabrication (Lewis and Nemme, 1997), Laser Engineered Net Shaping (LENS) (Griffith et al, 1997) and Electrophotographic Solid Freeform Fabrication (Kumar and Zhang, 1999).Traditional solid modeling systems do not support the ability to model, analyze or design heterogeneous components. Recently, methods for modeling such components have been explored. Kumar and Dutta (1998) have presented an extension of the traditional solid model representation based on regular sets (r-sets), to define rm-sets that include material data.

Kumar and Dutta (1998) have presented an extension of the traditional solid model representation based on regular sets (r-sets), to define rm-sets that include material data. They have defined set operations for combining these rm-sets to define heterogeneous solids. Jackson et al (1998) present an alternate representation where they subdivide a solid model into simpler domains over which composition functions are defined using Bernstein polynomials as basis functions.In this paper, a representation of composition distribution is presented that is based on a finite element mesh generated over the solid. The composition values at the nodes are interpolated to obtain composition within each element. In section 2, the method for shape and composition representation described. Elements that can be used for both modeling heterogeneous components and analyzing it using the finite element method are428described in section.

Elements that can be used for both modeling heterogeneous components and analyzing it using the finite element method are428described in section. A method for designing these components by optimizing the composition distribution is considered in section 3. A few examples are presented in section 4, where the design of components with varying composition and porosity are considered. Finally, conclusions are presented in section 5.2. Three-dimensional elements for composition representationThe material composition within a heterogeneous object can vary from point to point in an arbitrary fashion. To represent varying material properties within a solid, we propose a method that relies on a finite element mesh obtained by decomposing the solid into simple elements such as hexahedrons or tetrahedrons. The composition within these elements is obtained by interpolating the values of composition at the nodes of these elements using Lagrange basis functions.

The composition within these elements is obtained by interpolating the values of composition at the nodes of these elements using Lagrange basis functions. Thus by assigning appropriate composition values for the nodes of the finite element mesh we can control the composition within the solid. This method enables arbitrarily complex composition distributions to be specified provided that a sufficiently dense mesh of elements is used. Furthermore, the mesh used to represent the composition distribution can also be used to perform finite element analysis to obtain the mechanical properties of the structure. This approach was introduced in an earlier paper (Kumar and Wood, 1999) where two-dimensional elements were used to illustrate the idea.

This approach was introduced in an earlier paper (Kumar and Wood, 1999) where two-dimensional elements were used to illustrate the idea. Here we have described two three-dimensional elements that can be used to model composition distribution in a solid.Composition distribution is a volumetric property that can be expressed as a scalar or a vector field defined over the volume of the solid. A porous material can be described by defining its density distribution as a scalar field f =f (x,y,z). If the material is a composite of 'n' component materials, a vector of length n can be used to express the composition at any point. The first n-1 members of the vector can be used to represent the volume fraction of the first n-1 materials that make up the composite and the nth member can represent the void fraction (or density).

The first n-1 members of the vector can be used to represent the volume fraction of the first n-1 materials that make up the composite and the nth member can represent the void fraction (or density). Additional information may be required for some heterogeneous materials to specify other material properties such as the directions of anisotropy and fiber orientation of composites. All such volumetric properties may be expressed as a vector field v = v(x,y,z). Within each element volumetric properties can be expressed as follows by interpolating the values at the nodes of the element.v(r,s,t)=nvitsrN ,( ),ii= 1where, vi is the volumetric property vector at the node i, Ni are the basis functions used to interpolate the properties, n is the total number of nodes in the element. In this paper, we have used iso-parametric elements (Bathe, 1996) where each element is a perfect cube in the parametric coordinates.

In this paper, we have used iso-parametric elements (Bathe, 1996) where each element is a perfect cube in the parametric coordinates. The parameters r, s and t vary from -1 to 1 within each element so that each element is a cube of side length 2 in the parametric space. The basis functions mentioned above are also used to define a mapping between the parametric coordinates and the real coordinates x, y and z as defined below.x((cid:229)r,s,t)=ntsrNx ,( i i,),y(r,s,t(cid:229))=ntsrNy ,( i i),andz(r,s,t)=ntsrNz ,( ), i ii= 1i=1i=1429(2.1)(2.2)In the above equations (xi, yi, zi) are the coordinates of the node i. A variety of elements and their basis functions can be defined that can be used to represent the distribution of volumetric properties within a solid. As the number of nodes per element increases the degree of the basis functions used for the interpolation also increases.

As the number of nodes per element increases the degree of the basis functions used for the interpolation also increases. Two such elements and their basis functions are defined below.43t12s11 43t101312192s87r568167r1715 1856 14(a) 8-node element(b) 18-node elementFigure 1: 3D hexahedral elementsThe basis functions of an 8-node hexahedral element that is commonly used in finite element analysis is defined below. The element is illustrated in figure 1(a) and its eight basis functions can be expressed as:Ni=1 81(+rr i1)(+ss i1)(+tt i), i = 1,…8where (ri, si, ti) are the parametric coordinates of the node i. The basis functions and therefore the interpolation within each element are bilinear. Using this type of element, one can define simple piece-wise linear distribution of composition and material properties over a solid.

Using this type of element, one can define simple piece-wise linear distribution of composition and material properties over a solid. For example, we may want to specify that a solid of material A has a coating of material B such that its composition varies linearly from B to A within a specified depth from the boundary of the solid. This can be achieved by generating a mesh of 8-node hexahedral elements such that the first layer of elements on the boundary is of the appropriate thickness so that a linear variation from pure B to pure A can be represented. Figure 2 illustrates two hexahedral elements with nodal compositions defined such that the composition varies linearly from a red material at the top to a blue material at the bottom.(a) Mesh(b) Composition distributionFigure 2: Linear variation of compositionMore complex distributions require higher order elements. An eighteen-node element can be defined that allows a cylindrical variation of composition.

The nodal values are defined such that x = 0 for nodes 1-8, x = 0.5 at all other nodes except 13 and 18 where x = 1.0 (figure 1(b) shows the node numbering used here).(a) Mesh(b) Composition distributionFigure 3: Cylindrical composition distributionAs mentioned earlier, the finite element mesh used to represent the composition distribution can also be used to perform a finite element analysis to obtain structural and other mechanical properties of the solid. Since the composition varies from point to point within the solid so does the material properties such as elastic constants and density. An appropriate relationship between material properties and the composition must be established in order to compute the material properties at each point. This relation, which we will refer to as the material property function, depends on the microstructure and may vary if the component is made using different techniques even if the composition is identical.

The true relation between material properties and composition for a composite material can perhaps be established only by experimental studies. Empirical relations may be established by fitting a polynomial over experimentally obtained results. In this paper, we assume the material property functions to be simple polynomials to illustrate how mechanical properties may be computed if the material property function is known.For a linear elastic non-porous composite material with just two component materials A and B, we assume that the Young's modulus is a linear function of the volume fraction of material A.

Therefore, the relation between the Young's modulus of the composite material and the volume fraction of A can be expressed asEx=E AA+ 1(A)EB,where, x A is the volume fraction of the material A while EA and EB are the Young'smodulus of the materials A and B respectively.431(2.4)(2.5)If the material is porous, it can be thought of as a composite material that is made of material and void. The composition can be expressed using a relative-density field. In this case we assume the material property to be a polynomial function of density. Assuming the material to be linear elastic, we can express the Young's modulus as,nEE (2.6) where, E0 is the Young's modulus of the fully dense material, n is an integer and x is the relative-density or volume fraction of material.

This type of material property-density relationship has been used in many previous work on shape and topology optimization where the corresponding material has been referred to as Solid Isotropic Material with Penalization (or SIMP) (Bends˘ e 1995, Zhou and Rozvany 1991, Kumar 1993).=x,0During the finite element analysis, we integrate the principle of virtual work over each element to obtain the element stiffness matrix and then assemble these together to obtain the global stiffness matrix. The deformation of the solid due to applied loads is represented by a displacement vector field that is interpolated using the same interpolation functions that is used to represent the composition distribution within each element. This representation yields the following strain-displacement relation for the element.e }u]{ he [}{B=e(2.7) }{ e hu is the nodal displacement vector corresponding to the finite element.

This representation yields the following strain-displacement relation for the element.e }u]{ he [}{B=e(2.7) }{ e hu is the nodal displacement vector corresponding to the finite element. The governing equations can be expressed in general using the following type of weak form of a variational principal. XBDBX } dV,where,(cid:242)(cid:242)(cid:242)d+d=dT bX }{}T fX }{}T []T [}{{dS][{]{dVV The above equation represents the weak form applied to an element. V is the volume of the element, S is the surface area of the element that lies on the boundary of the object, {f} is the traction acting on these boundaries and {b} is the body force acting on the element. The vector {X} contains the nodal variables while the {d X} vector contains the corresponding virtual variables. The [D] matrix is a function of the material properties and therefore is a function of composition.

The [D] matrix is a function of the material properties and therefore is a function of composition. Since composition varies from point to point this matrix is also a function of the spatial coordinates. The left-hand side of equation (2.8) needs to be integrated to determine the stiffness matrix. The integration was performed using Gauss quadrature algorithm (Bathe, 1996).SV3. Optimization of composition distributionThe mechanical properties of a heterogeneous component can be determined by finite element analysis as described in the previous section. Alternately, one may want to design a heterogeneous component that has the desired mechanical properties. For example, one may solve for the optimal composition distribution with the objective of maximizing the stiffness of a structural component subject to a limit on its weight or mass. Maximizing the stiffness of a structure is equivalent to minimizing its compliance.

Maximizing the stiffness of a structure is equivalent to minimizing its compliance. The compliance is computed as a product of applied force times the deflection. It is a commonly used objective function in many structural optimization studies and has been used extensively in topology optimization research (Bendsoe and Kikuchi, 1988, Kumar and Gossard, 1996). Below we illustrate the process by designing a simple heterogeneous432(2.8)solid by minimizing its compliance. The solid is represented using the 8-node hexahedral elements described earlier. The material is assumed to be a two-component composite and the volume fraction of the primary component, denoted by x , is the design variable that is optimized.

The material is assumed to be a two-component composite and the volume fraction of the primary component, denoted by x , is the design variable that is optimized. The optimization problem may be written as, +W d(cid:242)(cid:242)x(cid:215)x(cid:215)x=GMinimize))ufd)(u ((utL)(subject to, x=x V d )((cid:242)V 0,{ } [ Dedtx)(]{ } de0=L(u)0£x£(3.4) 0 L(u), the mean compliance, is twice the work done by the applied forces (traction t and body force f) during the displacement u. Equation (3.2) describes the constraint that the total volume V of the primary material should be less than or equal to V0. {e } and {d e } are the strain and virtual strain in the structure caused by the displacement u and the virtual displacement d u respectively. [D(x )] is the matrix of elasticity constants that relate stresses and strains for a linear elastic material.

[D(x )] is the matrix of elasticity constants that relate stresses and strains for a linear elastic material. We assume that these elasticity constants are functions of the composition of the material.1The optimization problem given by the equations (3.1) - (3.4) can be solved using a nonlinear programming algorithm. Such optimization algorithms modify the design variables (composition vector at the nodes) in an iterative fashion searching for the optimal distribution. The objective function value (or compliance) is computed using the finite element method at least once during each iteration. Since each objective function evaluation is expensive, an optimization algorithm that does not require excessive objective function evaluations is desired. In addition, due to the large number of nodes in the mesh used to represent the composition distribution we need algorithms that can handle large number of variables.

In addition, due to the large number of nodes in the mesh used to represent the composition distribution we need algorithms that can handle large number of variables. Our work uses a variation of sequential linear programming, which is described in Kumar (1993).4. Examples and ResultsTwo examples of designing heterogeneous parts are presented here. In the first example, the part is assumed to be made of a two-material composite structure whose geometry is fixed but its composition can be varied to maximize its stiffness subject to a constraint on the final percentage volume of the primary material that is used. Figure 4(a) shows the geometry of the part, the applied loads and boundary conditions. The blue arrows represents normal pressure applied to the surface and the structure is supported at the back so that all the nodes on the back face are fixed (denoted using the yellow triangles).

The blue arrows represents normal pressure applied to the surface and the structure is supported at the back so that all the nodes on the back face are fixed (denoted using the yellow triangles). The Young's modulus of the primary material (Ep) was assumed to be 2x1011 units and Poisson's ratio of 0.3. The Young's modulus of the secondary material is assumed to be 0.5Ep. The material property function was assumed to vary linearly with composition as in equation (2.5).

The Young's modulus of the secondary material is assumed to be 0.5Ep. The material property function was assumed to vary linearly with composition as in equation (2.5). Figure 4(b) shows the optimal composition distribution that was computed using the constraint that the total volume of primary material (black in color) in the final part should less than or equal to 50% the total volume of the structure.433(3.1)(3.2)(3.3)a) Mesh, Load and Boundary conditionsb) Optimal composition distributionFigure 4: Optimal composition distribution for a cantileverThe composition variation within the part can be clearly visualized only by plotting the composition distribution on various section planes that go through the solid.

In figure 5, various section planes are plotted superimposed on a semi-transparent image of the part.Figure 5: Composition distribution on various section planes434Another example of composition variation was studied where the geometry of the part considered in the previous example is optimized by assuming that the material is porous so that the two materials in the composite are the primary material (Ep = 2x1011 units) and void. The volume fraction of the material, x , (or relative density) is optimized such that the overall volume is only 50% of the original volume of the part. This approach for geometry optimization has been used in many previous studies on shape and topology optimization. Regions where x =1 is fully dense and the regions where x =0 has no material. The mesh, the applied loads and boundary conditions shown in figure 4(a) was used in this example also.

Regions where x =1 is fully dense and the regions where x =0 has no material. The mesh, the applied loads and boundary conditions shown in figure 4(a) was used in this example also. The material property function was assumed to be quadratic as in equation (2.6) when n = 2.(a) Viewed from front(b) Viewed from behindFigure 6: Optimal geometryThe optimal geometry that was computed is displayed in figure 6, where regions with low value of x were made transparent. The transparency was varied such that the material is fully opaque when x =1 and fully transparent when x =0. The color is also varied such that the material is black in regions where x =1 and white where x =0 and various shades of gray in between.5. ConclusionIn this paper, a mathematical representation for shape and composition distribution using a finite element mesh is described. Two hexahedral elements and their interpolation functions are described that can be used to model heterogeneous solids.

Two hexahedral elements and their interpolation functions are described that can be used to model heterogeneous solids. The variation of the composition within each solid can be controlled interactively by modifying nodal density values or can be computed automatically by optimizing the composition distribution to meet design objectives and constraints.6. AcknowledgementsThe authors would like to acknowledge assistance from Mr. Lichao Yu and Mr. Aaron Wood in implementing the finite element analysis software. Partial funding from ONR contract N00014-98-1-0694 and NSF contract DMI-9875445 is gratefully acknowledged.7. 8. References [1] Bathe K. J., 1996, Finite element Procedures, Prentice Hall.435[2] Bends˘ e, M. P. and Kikuchi, N., 1988, “Generating optimal topologies in structural design using a homogenization method”, Computer Methods in Applied Mechanics and Engineering, Vol. 71, pp. 197-224.[3] Bends˘ e, M.

SIGNING CONFORMAL COOLING CHANNELS FOR TOOLINGXiaorong Xu, Emanuel Sachs, Samuel Allen, Michael CimaMIT Cambridge, MA 02139Contact: Emanuel Sachs MIT (617) 253-5381 sachs@mit.edu 77 Mass Ave Cambridge, MA 02139 Rtn 35-134Abstractto produce tooling with cooling channels which are conformal to the molding cavity. 3D Printed tools with conformal cooling channels have demonstrated simultaneous improvements in production rate and part quality as compared with conventional production tools. Conformal Cooling lines of high performance and high complexity can be created, thus presenting a challenge to the tooling designer. This paper presents a systematic, modular, approach to the design of conformal cooling channels. Recognizing that the cooling is local to the surface of the tool, the tool is divided up into geometric regions and a channel system is designed for each region.

Recognizing that the cooling is local to the surface of the tool, the tool is divided up into geometric regions and a channel system is designed for each region. Each channel system is itself modeled as composed of cooling elements, typically the region spanned by two channels. Six criteria are applied including; a transient heat transfer condition which dictates a maximum distance from mold surface to cooling channel, considerations of pressure and temperature drop along the flow channel and considerations of strength of the mold. These criteria are treated as constraints and successful designs are sought which define windows bounded by these constraints. The methodology is demonstrated in application to a complex core and cavity for injection molding.SFF technologies have demonstrated the potential1.

The methodology is demonstrated in application to a complex core and cavity for injection molding.SFF technologies have demonstrated the potential1. Introduction1.1 MotivationThe cooling of injection molding tooling is crucial to the performance of the tooling influencing both the rate of the process and the resulting quality of the parts produced. However, cooling line design and fabrication has been confined to relatively simple configurations primarily due to the limits of the fabrication methods used to make tools, but also due to the lack of a design methodology appropriate for cooling lines.131For many years mold designers have been struggling for the improvement of the cooling system performance, despite the fact that the cooling system complexity is physically limited by the fabrication capability of the conventional tooling methods.

Different methods such as the helical channels, the baffled hole system, the spiral plug system and the heat-pipes have been developed for the uniform and efficient cooling of the part (2). Some mold manufacturers such as Innova built tools with conformal cooling effect by stacking slices of the tool layer by layer with cooling channels milled on each layer. Other manufacturers such as CITO Products Inc. developed the pulse cooling technique for better control of the mold temperature in order to reduce the energy consumption and enhance uniform cooling condition.The emergence of Solid Freeform Fabrication (SFF) processes offers injection mold manufacturers new degrees of freedom in mold tooling. The SFF processes are additive processes that construct 3D objects by incrementally building up cross sectional layers of arbitrary complex shapes converted from CAD models.

The SFF processes are additive processes that construct 3D objects by incrementally building up cross sectional layers of arbitrary complex shapes converted from CAD models. Typical SFF processes include three dimensional printing, solid ground curing, stereolithography, selective laser sintering, laminated object manufacturing, etc. [23]. The ability to fabricate 3D features with almost arbitrary complexity makes these processes extremely useful for fabricating parts and tools that cannot be practically made by other techniques. One example of the applications is to fabricate complex cooling channels inside injection molds in order to improve the uniformity of cooling. The 3D Printing Lab at MIT has been participating in the injection molding tooling project for several years [24, 25, 26].

The 3D Printing Lab at MIT has been participating in the injection molding tooling project for several years [24, 25, 26]. The industrial application of 3D Printed tools with serpentine conformal cooling channels built inside have achieved the simultaneous improvement of the cycle time by 15% and the part distortion by 9% [26]. With the manufacturing flexibility offered by SFF processes such as 3D Printing, more complex cooling channel systems such as that shown in Figure 1b can be fabricated for the further improvement of the cooling performance.The emergence of new processes that can be used to create tools with conformal cooling channels placed with almost arbitrary complexity not only offers the designer new degrees of freedom in the design of injection molding tools but also simplifies the methodology used to design cooling channels.

The work discussed in this paper seeks to develop a methodology for the design of cooling channels that both simplifies the design and results in substantially improved performance. As you will see from the paper this methodology makes the design of complex cooling channel system for the inserts such as that shown in Figure 1b a handy work.132Figure 1. Left: solid model of the core insert with conformal cooling channels built inside. Right: solid model for the conformal cooling channel set1.2 Related WorkBefore the scientific analysis was introduced into injection molding, the design of the mold cooling system was dominated by designers' experience and simple formula [1-4]. It is not until early 80's that the mold cooling simulation was paid more and more attention. Different methods were proposed to predict the temperature field for the mold and the part during the cooling stage [5-14].

Different methods were proposed to predict the temperature field for the mold and the part during the cooling stage [5-14]. The cooling related quality issues such as residual stress, shrinkage and warpage were also addressed in [15-22]. Among those simulation algorithms, an iterative hybrid scheme proposed by Cornell Injection Molding Program (CIMP) became a standard scheme for mold cooling analysis due to its computational efficiency [8-12]. Today most of the mold design packages such as C-MOLD are equipped with this analysis scheme. This scheme treats the plastic part as one-dimensional transient heat transfer and the mold as three-dimensional heat transfer. The periodic transient mold temperature field within an injection cycle is separated into The quasi-steady component a quasi-steady component and a time-varying component. reflecting the cycle-averaged temperature field is obtained by solving the Laplace equation for the entire mold using the boundary element method.

reflecting the cycle-averaged temperature field is obtained by solving the Laplace equation for the entire mold using the boundary element method. The solution is then used as the boundary for 1D part temperature field. The iterative reference between the boundary element solution of the cycle-averaged mold temperature field and the finite difference solution of the part temperature field continues until a steady temperature boundary is achieved at the mold-part interface. Compared with the finite element simulation that needs to calculate the internal nodes, the hybrid scheme discussed above significantly reduces the computational time because only the boundary nodes are considered for obtaining the cycle-averaged mold temperature field.The emerging techniques for freeform fabrication of the injection molding tools place a new challenge to the mold design and analysis strategy due to the increased complexity in cooling channel geometry.

This situation motivates us to develop a systematic tool for conformal cooling line design. Compared with the existing cooling analysis software, the methodology133discussed ill this .paper builds .acisynthesis·tool instead of an analysis ••tool for the design of the complex conformal coolingcharuiels which take the full advantage ofSFF processes.2. Heat transfer model for conformal coolingA.~the~a1t1eiITlPlies,¢onformalcoolingis usedto signify cooling channels that conform toth~sllr~(lcYieft~~ meldigaXit;y. How~ver, in this paper,. th~ tel'fll"conformal cooling" has a further/~~pn.ifjc(ll1~y~atis(r~lated to the transient heat transf~r within the mold. When a mold is started\1gitit~essoIl1etime .before the mold reaches •• a·steady state· operating temperature. Figure2s~owsthem91c1.surface temperatur~historiesrecorded by the thermocouples for both themol4',Vith conf0srn-aLcooling channels and that with straight cooling channels [24].

Figure2s~owsthem91c1.surface temperatur~historiesrecorded by the thermocouples for both themol4',Vith conf0srn-aLcooling channels and that with straight cooling channels [24]. As one can see~omthefigllre,the mold surface temperature of the core with straight cooling channels trackYd.(ov~r25sucgessiveinjections startingfromthe coolant temperature of 12° C and reaching a cycle average steady state temperature of apgroximately 55° C. However, if the cooling lines are placedyery close.andconformaltothe. mold surface the steady state. condition is reached very quickly. As illustrated in Figure 2 the cycle. average temperature of aconformally cooled core reaches its steady state value after one injection cycle. Our operational definition of 90nfonnaLcooling then isthatthe cycle ayerage temperature reaches its steady state value within one. inj~9tioncycle.

As the hot .plasticieomesin during each sU9cessiv~injection, heat transfer take~ p~a~e acros~ the plastic ll10ldinterface and a heat pulse is condu.cted through the mold matenalltself. ThIS heat pulse warms up the mold Il1aterial as it propagates toward the. cooling channels and is eventually134removed in the cooling water. If the cooling channels are far from the mold's surface, successive heat pulses keep raising the temperature of the rnolduntil the heat. pulse propagating in is balanced by heat extraction by the coolant. Ifthe. cooling channels are close to tbemold surface, the effective thermal mass of the tool is confined to thatregion <between< the surface and the In addition, the conduction path from the surface ofthe cooling channels and is much reduced. tools to the cooling lines is reduced.

tools to the cooling lines is reduced. As a result, the steady state condition is reached much more rapidly and can inTact be attained within one injection cycle.An energy balance may be written for the active portion of the mold, that is the portion between the surface and the cooling lines. Equation (1) shows the resulting differential equation (see Appendix A for derivation of this equation).C I dTm +Pm m m dthrcDKm 2KmW + hrcDlm(T -T) cmP pCpIp (Tme1t t cycleTe.iect)(1)The first term in Equation (1) captures the thermal mass of the tool and the build-up of heat as the temperature of the tool increases. The second term in Equation (1) captures the transfer of heat by conduction through the mold and then convection into the cooling fluid. The right hand side of Equation (1) captures the source of the heat, which is the cooling down of the plastic. This first order differential ordinary differential.

The right hand side of Equation (1) captures the source of the heat, which is the cooling down of the plastic. This first order differential ordinary differential. equation has the solution of the form shown in Equation (2) where Tms is the cycle averaged mold temperature at steady state and 1: is the time constant of the system. Equations (3) and (4) give the expressions for cycle averaged mold temperature and the time constant respectively. Our definition of conformal cooling can now be stated formally by requiring that 1: be less than or equal to one injection cycle time. Figure 3 shows a prediction of Equation (2) superimposed onthe experimental results previously shown in Figure 2.

Figure 3 shows a prediction of Equation (2) superimposed onthe experimental results previously shown in Figure 2. As can be seen, there is reasonably good prediction with the cycle average temperature.(2)(3)(4)A limiting case of Equation (4) is that where the heat transfer to the fluid is very efficient and we can then examine the limiting case where the heattransfer coefficient goes to infiliity.In this case, the expression for the time constant reduces tothe form showninEquation (5). Inthis simplified expression we see that the important material property for the mold is the thermal diffusivity Which.is ·Km / PmCm. We also. see that the time constantjsproportionaltot~esqua~e of the distance between the surface of the mold and the cooling channels. This. simplified expression makes ciear the importance of consideringthisasatransientiheattransfer calculation.

This. simplified expression makes ciear the importance of consideringthisasatransientiheattransfer calculation. If this were a steady state heat transfer problem then doubling the thermal conductivity of the135mold would allow the channels to be placed twice as far away. However, as we can see from Equation (5) if we double the thermal conductivity and place the channels twice as far away the time constant in fact increases by a factor of 2. Thus, while the material properties are important, the geometry (as seen by the square of the distance of Equation (5)) is even more important.= PmCml~ K m(5)2520',\ '1 ' \:>".) .'.\ : I \ .,\::\: ,. , \ I. . .\;~., )\ , , , " ; .'I,~\':' ; , . • '.' .~\ : .\ "I ~\;, " ol • J : .~ .. '"'. " ".',\: ~, \t, , " I:'\ . . ,\I,, \, ,,'I~,I','\.. , " .' \;), " "."\'_sirrulation data ....... experimental dataTime [sec]Figure 3. ComparisoDoftheexperimen(1atavs.

. ,\I,, \, ,,'I~,I','\.. , " .' \;), " "."\'_sirrulation data ....... experimental dataTime [sec]Figure 3. ComparisoDoftheexperimen(1atavs. the simulation data for the mold surface temperature profile during successive. injections3. Design methodologyThe introduction of· conformal cooling significantly simplifies the .injection molding cooling system design methodology. In theconfonnal cooling situation, the· heat transfer is localized in a small region between. two adjacent cooling channels. This feature suggests that we first design a cooling "cell" composed ofthe small region between the adjacent cooling lines and thenmap the solution to the entire mold. The flexibility of SFF pro.cesses makes this modular approach possible by minimizing the manufacturing constraint that must be applied. This strategy simplifies the cooling line design by providing a sequential approach which provides a global solution by the addition of many local solutions..

This strategy simplifies the cooling line design by providing a sequential approach which provides a global solution by the addition of many local solutions.. While the design process is simplified, the resulting cooJing line designs can be quite complex and take full advantage of the flexibility of SFFprocesses.Figure 4 illustrates this design strategy by using a generic·part with a hemispherical dome alld a flat bottom. AsSl10wn inthe figure, the. part is first divided into two cooling ZOl1eS (a hemisl'here<alld a flat.surface) based on its geometry. Then in each coolil1g zone the conformal cooling surface is const11lcted~ndthe cooling channel topological structure is defined. After that the system of cooling channels is further decomposed into small elements called cooling cells.136The heat transfer analysis and the cooling system design is based on these cooling cells and is then mapped to the entire mold.

This modularized design strategy is not sensitive to the part therefore it keeps the same design simplicity no matter how complex the part geometry, geometry is.conformalparcoolingcavity!0 0 00o oCooling zone I,back to moldIContormal cooling design rulesOriginal\Cooling cell beforeCooling cell afterMap design back to moldCooling zol1e2Figure 4 Steps for the modularized cooling line design for a generic part4. Design rulesAfter the cooling systemha.s pee~ .• dec0tnPosedjnto>simplecooliIlg cells .by the method discussed above,. the design rules are applied. to these cooling cells .• in order toobt~ilJ.cooling channel design parameters and process conditions. In thischapt~r, six.design·nl1esaryproposed and design windows. are constructed for the cooling line design based on individual cooling cells. These rules include design for conformaLcoolingcondition,<designJorcoolant pressure drop, design for.

are constructed for the cooling line design based on individual cooling cells. These rules include design for conformaLcoolingcondition,<designJorcoolant pressure drop, design for. coolant temperature uniformity,designJor .sufficient .cooling, design for. uniform cooling and design for mold strength and deflection.4;1 Design for conformal cooling conditionThe conformal cooling condition defined in section 2 must be applied throughout the tooling in order to guarantee the good control of the temperature at the surface of the tool. In order to satisfy this condition, the mold designer can increase the heat transfer coefficient, increase the channel diameter, decrease the distance between the cooling lines and the mold wall137or choos.e mold material with a high thermal diffusivity.

The determination of these design parameters is also constrained by other design rules that will be discussed below.4.2l)esigp for coolant pres~ured.ropis constrained by the available pumping pressure of the chiller.

The objective of the cooling line design for pressure drop is to Jind a proper colllbination ofthe coolant flow rate, the cooling chall11el dianletElr ap.gthecp()ling line length so that the resulting total pressure drop is smaller than the given press.ure~udgElLThe fluid mechanics of the incompressible flow can be used to predict the coolant pressure drop that is a function of the cooling line length, the cooling line diameter and the coolant flow rate [27].The allowable pressure drop of the coolantin the.conformal cooling channelL P=-pv Cf 2D2(6)where Cf in Equation (6) is the cooling channel surface friction factor which differs for different flow regions:(for laminar flow)(7)C = 0.25(10 .• ·..[f-!-)I.11+ 6.9])-2 gro "3.7D f~D1.8 2(for turbulent flow)(8)In the above equations, ReD is the Reynolds number of the coolant flow.

e is the surfaceroughness of cooling channels.4.3 Design for coolant temperature uniformityThe objective of the design for the coolant temperature uniformity is to check and make sure that the coolant temperature drop is maintained within a certain range. A simple expression of the coolant temperature drop LIT is obtained by the following equation:(9)where lp is half the plastic part thickness, w is the cooling line pitch distance, L is the cooling line length, Q is the coolant flow rate and tcycle is the injection cycle time. pc, Cc, PP' Cp are the densities and specific heats for coolant and part materials respectively. Equation (9) indicates that during the steady injection cycles the heat pulse due to the cooling down of the plastic part is totally converted to the temperature rise of the coolant flow.

Equation (9) indicates that during the steady injection cycles the heat pulse due to the cooling down of the plastic part is totally converted to the temperature rise of the coolant flow. In order to reduce the coolant temperature drop, the designer can use the coolant with large thermal mass, increase the coolant flow rate, decrease the pitch distance between two adjacent cooling channels or reduce the length of the cooling line.1384.4 Design for sufficient coolingAs we have discussed in section 1, the cooling analysis scheme adopted by mostof mold design software is computationally expensive and not good for the design and analysis of complex cooling channels. With the concept of conformal· cooling, this scheme is much simplified. As the matter of fact, the steady cycle averaged mold temperature can be directly derived from Equation (3). However, this expression requires both the part ejection temperature Teject and the cycle time tcycle which can not be both obtained.

However, this expression requires both the part ejection temperature Teject and the cycle time tcycle which can not be both obtained. A simple iteration discussed below the cycle time and the steady cyclic mold temperature T ms based on the required part finds ejection temperature:Step 1. Assume the cycle averaged mold temperature Tms . Step 2. Calculate the cycle time tcycle for the required part ejection temperature according topart heat transfer.Step 3. Calculate the part ejection temperature at the end oftcycle. Step 4. Calculate the cycle averaged mold temperature Tms based on equation (3). Step 5. Replace Tmsin step 1 by the cycle averaged mold temperature value obtained instep 4. Step 6.

Step 4. Calculate the cycle averaged mold temperature Tms based on equation (3). Step 5. Replace Tmsin step 1 by the cycle averaged mold temperature value obtained instep 4. Step 6. Follow the iterations reaches a steady value.from step 1 to 5 until the cycle averaged mold temperature TmsFor readers with special interest in detail algorithms, please refer to ApPcendix B.4.5 Design for uniform coolingThe "uniform cooling" in this paper has both the globaFand the JocaLmeanings. The global uniformity is the cooling rate variation over the entire mold. It is guaranteed by/keeping the coolant temperature uniformity. The local cooling uniformity refers to the variation. of the mold surface temperature within the individual cooling cell sketched in Figure 5(a). The local cooling uniformity is defined by the difference of the cycle averaged temperatures onthe mold surface right above the cooling channel and at the middle oftwo adjacent channels, i.e.

The heat conduction distances la and lbin equation (19) and (20}>ary.ch()s~~tobedistancesfrom point A and B. to the wall of the coolin? (channel respectively. As one can .see from the figure, the analytical and numerical solutions match very well in a certain pitch distance range. More accurate prediction can be achieved by adding adjustment factorst? la and h,4.6 Design for mold strength and deflectionR.ao [3] pred.icted the mold stress and delection based on the rectangll,lar cooling channel model shown in Figure 6. According to his model, the maximum tensile stress in the mold under a certain injection pressure Pm is:O"max =0. 5PmD2 12 m(13)The maximum shear stress in the mold is:Tmax =0.75PmD I(14)mThe maximum :m.old surface deflection under pressurePm is:(15)The above expressions represent the worst case of the loading because the commonly used cooling channels are circular shaped that result in much smaller stress and deflection.

The numerical simulation shows that the stress concentration is reduced by over 50% if we choose channels with round comers.Inj ection pressure PmMoldMold surface profile under injection pressureCooling channelFigure 6 Sketch of a cooling ceHunder the injection pressure5. Tool design and fabrication with conformal cooling channelsWith the help of the .conformal cooling line design methodology, the 3D Printing Lab of MIT has successfully designed and fabricated the. moldinserts for the part shown in Figure 7. The challenge placed by this p~t't is thatthe geometric features are so<small andsoclos~toeach other that it gives the designer little room for the cooling channel placement. In order to proceed the conformal cooling line design for this part, we first divide the part into two cooling zones. For each cooling zone, the cooling line topological structureiisdefined and differentcooling channel cross section shapes are assigned to differenfsectionsofthe cooling line..

For each cooling zone, the cooling line topological structureiisdefined and differentcooling channel cross section shapes are assigned to differenfsectionsofthe cooling line.. Then the cooling system is further decomposed into individual coollngcellsand the design rules are applied to these cells to obtain the local solution of the cooling channel parameters. The 2D numerical simulation (see Appendix C for details) is also used for the evaluation of the transient heat transfer in each cooling cell..Figure 9 shows the design results for individuaLcooling cells. These results include the cooling channel geometry,processcondition as wellas the evaluation of design ruJes such as the part ejection temperature, the coolant pressure drop, the mold stress, etc. In addition to the local design of individual· cooling cells, the designer is responsible for coordinating between local solutions for the •. properglobal••• perf0rtnan.c¢.0l1e~xa.t1lpleiiofsucll coordination is to balance the.

The .stl file is then created for solid freeform fabrication. Figure 10 shows the resulting green part fabricated by 3D Printing process.Figure 9 Green part of theprocess based on6. ConclusionsSolid Freeform Fabrication processes such as 3D Printing can create injection molding tooling with complex cooling channels offering the potential for substantial improvement in production rate and part quality. This capability raises the.challenge of designing the cOmplex cooling channels required to realize these improvements.This work presents a systematic method for the design of cooling channels for tooling.First the mold surface is. decomposed into manageable sections called cooling zones. Then system of cooling channels is designed for each cooling zone. This system ofcooling channels is further decomposed into smaller elements called. coolingi cells that are easy to analy*e. The cooling cell is a sandwich.

This system ofcooling channels is further decomposed into smaller elements called. coolingi cells that are easy to analy*e. The cooling cell is a sandwich. structure covering the part,the mold and·.the cooling channel region between two adjacent cooling •lines. Six design rules are applied in order to createidesign windows for the individual cooling cell. These design rules include design for conformal cooling condition, design for coolant temperature drop, design for part ejection temperature, design for143sufficient cooling, design forffiold strength and deflection and design for cooling uniformity. After the design for individual cooling cell is fillished, the solution is mapped back to themold in order to build the entire conformal cooling sY$tem.7.J\ckno'Nledg~entThe work discussed in this paper is supported by the Technology Re-Investment Project, Cooperative Agreement (DMI - 9420964) funded by DARPA and administered by NSF, and the 3D Printing Industrial Consortium.

The authors thank members of3DP Consortium and the colleagues of 3DP Lab for their consistent support and encouragement. Special thanks are due to Dick Barlik from Hasbro, who designed the prototype part for the 3DP benchmark tool, David Brancazio and Jim Serdy from 3D Printing Lab of MIT, who took care of most of the work from administration to manufacturing of the conformal cooling project, Dr. K. K. Wang and Dr. Hieber from Cornell Injection Molding Program, who provided generous help and encouragement in mold cooling system design and simulation.8.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDesigning Functional Beauty through Additive Manufacturing: Prototyping ofRunning-Specific Prostheses Using Selective Laser SinteringShoichi Sato*, Naoyuki Togo* and Shunji Yamanaka*Prototyping and Design Laboratory,Institute of Industrial Science, The University of TokyoAbstractThe objective of this research is to establish a new methodology of designingaesthetically and functionally satisfying mass-customized products that fit individual bodies. Theprimal phase of this study, the prototyping of additive manufactured Running-Specific Prostheses(RSPs), is shown in this paper. The focus of this work is to present the capability ofmanufacturing such products by using Additive Manufacturing (AM), especially LaserSintering (LS).

The focus of this work is to present the capability ofmanufacturing such products by using Additive Manufacturing (AM), especially LaserSintering (LS). The first section describes the method in which a design process that uses AMtechnologies is established, aiming mainly to achieve a functional product that is alsoaesthetically pleasing. The latter section presents engineering verifications of previously designedprostheses, and the application of structural improvements to achieve enough reliability for usertests. The improved prosthesis was tested by a transtibial amputee runner; user review andremaining issues are reported.IntroductionAM technology has received great attention in recent years, however, its actual usage inthe industry is limited. It is usually used as a fundamental prototype for mass products (generallyreferred to as Rapid Prototyping) and to produce various types of jigs for large-scale factories(Rapid Tooling).

It is usually used as a fundamental prototype for mass products (generallyreferred to as Rapid Prototyping) and to produce various types of jigs for large-scale factories(Rapid Tooling). These types of applications can be beneficial due to the speed and simplicity ofthe AM process. Recently, there are more final products being manufactured with AM, such asaviation parts [1] and in the medical industry [2]. However, it is estimated that 90% of theproducts manufactured by AM are mainly prototypes, while less than 10% are final products [3].1749For that reason, it can be pointed out that the AM process is inferior in accuracy, speedand cost when comparing to existing manufacturing processes such as subtractive manufacturingand injection manufacturing. According to Niino [4], if the performance of each process is takeninto consideration, AM has more weak points than existing manufacturing processes.

According to Niino [4], if the performance of each process is takeninto consideration, AM has more weak points than existing manufacturing processes. Toaccelerate the popularization of AM technology, a technical improvement that resolves thesenegative aspects is desirable. On the other hand, showing new value-added products that makeuse of the AM technology benefits is crucial.As previously described, AM process’ speed and simplicity are clear advantages. Onthe other hand, its flexibility of manufacturing freeform designs with almost no limitation is morenotable. The AM process enables making objects that existing processes cannot manufacture,such as highly complex freeform designs, minutely detailed objects and multi-layered innerstructures.

The AM process enables making objects that existing processes cannot manufacture,such as highly complex freeform designs, minutely detailed objects and multi-layered innerstructures. Niino [4] calls this characteristic a Complexity, and points out that, in the future, thelikelihood of the AM market expanding depends on finding the applications that could maximizethe benefits of Complexity rather than its speed.Additionally, AM technology is more suitable for mass-customized productioncontrasting to existing production technologies. The hearing aid industry has already started toutilize AM and more than 15 million hearing aid units have been made [5].

The hearing aid industry has already started toutilize AM and more than 15 million hearing aid units have been made [5]. There is also amovement of manufacturing customized shoe soles: some companies have created supplyingsoles that are optimized for individual foot shape by using AM technology [6].Making practical use of these benefits, we aim to establish in this research amethodology of designing value-added products that fit individual bodies. This project isconsidered a first milestone for a bigger goal, as we wish to create RSPs by using plastic-basedLS technology. Designing and manufacturing prostheses is mainly for a minority of people.Therefore, existing manufacturing systems that presuppose mass production cannot fit itsindividual needs sufficiently. This circumstance can be the cause of the problems that aredescribed in the following chapter.

This circumstance can be the cause of the problems that aredescribed in the following chapter. As we stated previously, the AM process will sufficientlyadapt for less volume and individually customized production. Therefore, an AM based systemcould generate new value while solving existing problems. Although specific information aboutstructure of prosthetic legs will be given in the next chapter, in this study we mainly focus about1750the development of prosthetic sockets.The present situation of prosthetic legsThe structure of prosthetic legs that are generally used is shown in fig. 1. They can beroughly divided in three sections: the first is called the socket to which the amputee’s residuallimb is attached, the second is the foot part that provides contact to the ground and the third isanother connective part and the joints between them. In terms of RSPs, leaf springs take the placeof the foot parts and are usually made of CFRP (fig. 2).fig.

In terms of RSPs, leaf springs take the placeof the foot parts and are usually made of CFRP (fig. 2).fig. 1 The structure of prosthetic legs fig. 2 RSP structureAs mentioned above, prosthetic legs are based on a standardized module-assemblysystem. Although that allows the supplying of affordable prostheses by combining metal parts,their appearances end up coming out quite unrefined. Therefore, amputees normally hide theirprostheses under their clothes, a fact that has prevented normalization for the user. Meanwhile,sockets are completely custom-made products that depend on the prosthetist’s craftsmanship.Fully handmade processes are still very common nowadays - almost no digital manufacturingtools are used. This could end up providing products that have an unstable design quality incontrast with other industrial products.

This could end up providing products that have an unstable design quality incontrast with other industrial products. Additionally, the manufacturing time and costs would beincreased and would eventually slow down the spread of prosthetic legs in regions that do notsupport financial aids.1751The significance of applying AM for prosthetic legsAM technology can automatically manufacture complex freeform objects like socketsby using 3D data. Hence, if prosthetists could use CAD software, they would be able toexclusively separate their designing operation from traditional manufacturing process and theycould concentrate on improving its quality or use their extra time for designing another socket.Accordingly, while decreasing the total cost of making sockets, prosthetists could respond to agreater amount and a wider range of user demands.

The quality of the products would bemechanically stable at the same time.On the other hand, AM enables to partly integrate the socket, the connective parts andjoints into one part. While being fit for the amputee’s individual set up, designing these integratedsockets that are both functionally and aesthetically pleasant is technically possible. Moreover, byusing the freeform characteristics of AM, there is a possibility of realizing a superiorperformance in strength and lightweight properties.Notably, we decided RSPs are an obvious example indicating possibilities that arementioned above. Comparing to regular prosthetic legs for daily use, RSPs need more strictstrength reliability and must be more lightweight. Additionally, as sport equipment, awell-designed appearance could give a mentally positive influence for audiences and athletesthemselves.

Additionally, as sport equipment, awell-designed appearance could give a mentally positive influence for audiences and athletesthemselves. Beyond the traditional values of prostheses - such as reproducing missing body parts– we also suggest that it will be significant for both amputees and society to create beautiful andhigh-performance prostheses that can be exposed, rather than hidden.fig. 3 Sprinting with a RSP1752ObjectivesAs an early phase of this study, the process of primal prototyping of transtibialrunning-specific prostheses is shown in this paper. One transtibial amputee athlete participated inthis study. First, an Additive Manufactured RSP (AMed RSP) was developed for that athlete.After collecting sufficient knowledge, we will attempt to apply this to other athletes. The profileof the athlete is stated in table. 1 below. The athlete is capable of sprinting 100m in less than14seconds.

The profileof the athlete is stated in table. 1 below. The athlete is capable of sprinting 100m in less than14seconds. During sprinting, the maximum load that strains the prosthetic socket is higher thannormal; therefore, we decided this case is reasonable for a primal trial.table. 1 The profile of the athleteAge22GenderFemaleProsthesisLeft foot transtibial prosthetic legCompetitionTook part in Paralympic, IPC Athletics World Championships andothersThe existing RSP that is used by the athlete is composed by a CFRP-made socket, acarbon leaf spring Cheetah Xtreme category5 (Össur hf.), and titanium made internal pyramidconnecting adaptors. The appearance of the RSP is shown in fig. 4 underneath.fig. 4 The existing RSP that is used by the athlete1753In this study, 3D digitizer and CT scan were used to measure the form of whole RSPthat is referred above.

4 underneath.fig. 4 The existing RSP that is used by the athlete1753In this study, 3D digitizer and CT scan were used to measure the form of whole RSPthat is referred above. The inner surface form of the prosthetic socket that is attached to theresidual limb was extracted exclusively, and was used as a basis of design. Although we considerit a final goal to scan the human body directly by using digital equipment (such as a hand-held3D scanner), we used a different process because this is our first attempt.Establishing design conceptBefore establishing the design concept, the problems behind the design of existingRSPs were taken into consideration. It is clear that, although existing RSPs functionally enableamputees to run, the aesthetic appearance and the regard for user-centered design are inadequate.The shape of the prosthetic sockets is quite similar to the shape of the amputee’s residual limb, soit can give the user a somewhat painful reminder.

Additionally, the mechanical parts that areconnected to the socket do not help create a visual harmony between the RSP and the humanbody. Furthermore, the sharp edges of those components can sometimes injure the other foot.Considering these situations, it can be said that an industrial designer taking part in themanufacturing process and improving aesthetic value and usability can be very beneficial.On the other hand, both the functionality that the RSP should achieve and themanufacturing characteristics of plastic-based AM processes were taken into consideration.

Eachitem is categorized and listed as follows:Functionalities that the RSP should achieve:Structural strength that can withstand human maximum speed running.Being lightweight while also having enough strength, as described above.Connectable structure for existing standardized parts such as leaf springs.Manufacturing characteristics of plastic-based AM processes:At this point, the material is Nylon12.Almost no limitation in manufacturing a highly complex freeform design.A form that allows extra material powder to be completely removed is preferable.Taking the above list into consideration, we attempted to generate a design concept. Asa first step, numerous pictures were collected of existing products that were made by AMprocesses, of sports equipment, organic structures and so on.

Asa first step, numerous pictures were collected of existing products that were made by AMprocesses, of sports equipment, organic structures and so on. Among them, the structure of bones1754was especially striking since the behavior and functionality of the bone structure seemedreasonably suitable for the AM process.The bone structure is composed of bone tissue, cartilage, periosteum and bone marrow.The bone tissue forms most of the rigid part of the bones and is comprised of the cortical andcancellous bone. The cortical bone is the outer shell of the bone structure and consists of a highdensity of cells, being therefore harder and stiffer. The cancellous bone (fig. 5), also known astrabecular or spongy bone, is the inner side of the bone. It is made of porous tissue and has ahoneycomb-like network structure of plates and rods. Due to these characteristic structures, thebone is both relatively strong and sufficiently light. The bones of birds (fig.

Due to these characteristic structures, thebone is both relatively strong and sufficiently light. The bones of birds (fig. 6) are especiallystrong and lightweight in order to be suitable for flying, and their structure consists of numeroustrussed trabeculas in a hollow inner space.fig. 5 The cancellous bonefig. 6 The cross section of a bone of birds（http://m.inmagine.com/image-spl007312-Bone-tissue%20-SEM.html）（http://www3.famille.ne.jp/~ochi/kaisetsu-01/kouzou.jpg）These characteristic structures are generated as a result of the bone remodeling processdone by osteoclasts and osteoblasts. By responding to mechanical stress, the structure isremodeled to increase strength around regions where it has been highly stressed and, in theopposite case, decrease density in areas with less stress. This physiological effect is generallyknown as a Wolff’s law.Bone remodeling, as we stated above, can be seen as a kind of structural optimizationand it fits AM technology.

This physiological effect is generallyknown as a Wolff’s law.Bone remodeling, as we stated above, can be seen as a kind of structural optimizationand it fits AM technology. By putting material in an appropriate amount and form around where1755it’s needed, the bone can achieve effective strength and become more lightweight. In addition,AM technology is capable of manufacturing complex structures with almost no limitation, beingtherefore able to also include highly intricate structures inside. In the end, we decided to use bonestructures as a main design concept reference.As a result of the above researches, we named the new design concept “the trabecularstructure concept”. It can be described as follows: a large number of small beams extend fromaround the residual limb to the leaf spring connecting section. The beams compose a complexstructure as they cross and combine.

The beams compose a complexstructure as they cross and combine. The density of the structure increases where stress is appliedwhile it decreases in other parts in order to remain lightweight. The outer form of the trabecularstructure follows the smooth curved surface in order to match the appearance of the socket withthe human body shape. Additionally, the holes and cavities of the structure will help extract thesurplus material powder from the inside.Based on the design concept above, Yamanaka made specific design sketches. As seenin fig. 7, the appearance of the socket was defined by numerous sketches.fig.

As seenin fig. 7, the appearance of the socket was defined by numerous sketches.fig. 7 the specific design sketch1756Manufacturing the design concept modelA study model that presents the design concept was made on the early research phase.The purpose of the model is to share the exact image of the design concept and to enable thediscussion of improvements on the later phase of the research with the research members.Firstly, by making reference to the design sketch, the 3D model of the socket was made.The CAD software that we used was Rhinoceros 5 for MAC (Robert McNeel & Associates).Secondly, the model was manufactured by an LS type AM machine RaFaEL300 (Aspect inc.).The maximum work size of the machine is 300mm in both width and length and 370mm in height.The work size allows the socket to be manufactured without parting it.

The material powderASPEX-PA2 (Aspect inc.), which is mainly composed with Nylon 12, was used to manufacturethe socket.The manufactured socket is shown in fig. 8. In the several early attempts, the socketwas not able to attach to the residual limb because of form inaccuracy. After adjustments to themachine were done, we succeeded in manufacturing a socket that has sufficient accuracy.Afterward, a test fitting was done as seen in fig. 9. As we stated previously, although the socketdoesn't include fully complete functionality on this phase, it was possible to attach, stand up, andwalk.fig. 8 The design concept modelfig. 9 Test fitting(Photo by Yasushi Kato)1757On this phase, we additionally carried out the research outreach activities mainly toresearch about impressions of new design concept. We organized an exhibition “Designing Body Making Beautiful Prosthetics” during June 4th to 14th 2015, in a gallery of The University ofTokyo.

We organized an exhibition “Designing Body Making Beautiful Prosthetics” during June 4th to 14th 2015, in a gallery of The University ofTokyo. The design concept model and the research promotion movie were exhibited to public asseen in fig. 10. We interviewed visitors for their impressions and received advice from externaltechnical experts, which were of benefit to the research.fig. 10 Exhibition (Photo by Yasushi Kato)Determining the strength benchmarkAs a next step, we applied improvements to the socket so that it could achieve anadequate reliability for the running tests. The strength benchmark was estimated first. Accordingto the benchmark, several components of the socket were developed.Firstly, we asked the athlete to do a maximum speed sprinting while using an existingRSP so that we could take a slow motion movie of it. As a result of image analysis (fig. 11), itwas estimated that the carbon leaf spring was deformed downward approximately 96mm.fig.

As a result of image analysis (fig. 11), itwas estimated that the carbon leaf spring was deformed downward approximately 96mm.fig. 11 Image analyses1758Secondly, we measured the spring constant of the leaf spring so that we could be able toestimate the elastic force of it. A carbon leaf spring Cheetah Xtreme category5 (Össur hf.) wasused for the compression experiment (fig. 12).fig. 12fig. 13 Load vs stroke trace,The compression experimentX-axis refers stroke, Y-axis refers loadThe conclusion was that the spring was deformed downward 52mm as it was stressedwith 2.5kN load (fig. 13). Hence, the spring constant was,k=P/δ=4.81×10*4[N/m]As a result, it can be estimated that approximately 4.6kN of elastic force was receivedthe leaf spring while sprinting.

13). Hence, the spring constant was,k=P/δ=4.81×10*4[N/m]As a result, it can be estimated that approximately 4.6kN of elastic force was receivedthe leaf spring while sprinting. Due to the law of action-reaction, the prosthetic socket receivedthe same amount of load.According to the estimation, the strength benchmark was determined; the socket shouldresist 15kN (FOS of 3) of vertically downward structural load on the central axis of the socket.Structural improvementWhile we designed the design concept model, the outer trabecular structures and theinner structure of the socket were designed as we prioritized aesthetic quality rather than thefunctionality. As a next step, structural improvement is needed to surpass the strength benchmark.

As a next step, structural improvement is needed to surpass the strength benchmark. Modifications of the CAD data of the socket were repeated, making reference to Finite Element1759Method (FEM) analysis.Through the improvement, it appeared that the modifications of the inner structurehighly affected the total strength of the socket. AM process is superior in manufacturing highlycomplicated freeform objects; hence the modifications previously mentioned match thecharacteristics of AM technology. The two pieces of topology optimization software such asOptishape-TS (Quint Inc.) and solidThinking Inspire (Altair Engineering, Inc.) were utilized togenerate referential form models that would be effective for increasing structural strength.A cross section of the socket in its primal phase is shown in fig. 14. Modifications areapplied afterward mainly to increase the complexity of the inner structure (fig. 15-16).

14. Modifications areapplied afterward mainly to increase the complexity of the inner structure (fig. 15-16). The shapeof the outer trabecular structure was simultaneously redesigned. The corrected models wereproduced at a high frequency by utilizing an AM machine as a Rapid Prototyping toolfig. 14fig. 15fig. 16Primal phase structureMid phase structureCurrent structureDesigning the internal adaptorFor practical use, a firm connection between the socket and the leaf spring is crucial.On the design concept model phase, the leaf spring was attached to the socket with standardizedM6 bolts and nuts. There was a concern that the stress concentration would occur around thecavities that enable the nuts to be set in the socket. As a result of a compression strength1760experiment, it was noted that the connecting section of the socket that includes the cavities wasbroken by approximately 5kN of structural stress (fig. 17).fig.

17).fig. 17 Cracked cavitiesAccordingly, an alternative solution was needed. The solution should assume thefollowing:1. Maintain a firm connection between the socket and the leaf spring while sprinting.2. Transmit the structural load over a large area to reduce the stress concentration on a Nylon part.3. Preferably integrate the adapters into the socket without post-processing.4. Be based on the characteristics and the limitations of AM process.As a result of the design work, we designed the method that uses existing bolts andinternal adaptors that are made with metal. The internal adaptors are inserted into the back of thesocket and the leaf spring is attached to those adaptors by using standardized bolts (fig. 18).fig. 18 The Structure of RSP in case of using the internal adaptors1761The adaptors were manufactured by a subtractive process. 5052 aluminum alloy wasused to manufacture the adaptors (fig. 19).

18 The Structure of RSP in case of using the internal adaptors1761The adaptors were manufactured by a subtractive process. 5052 aluminum alloy wasused to manufacture the adaptors (fig. 19). The length of the adaptor is 25mm and the maximumdiameter is 14mm. There are male screw threads that are characterized by the specific screwprofile on the outer periphery of the adaptor. The screw profile (fig. 20) is composed of arcs andtangents to avoid stress concentration. On the central axis of the adaptor, there is a hole withinternal threads corresponding to a standardized M6 bolt.fig. 19fig. 20The internal adaptorThe internal adaptorDesigning the running test modelBy integrating the knowledge obtained previously, the socket was redesigned as fig. 21shows and the sketch was provided by Yamanaka. The major changes from the design conceptmodel are as follows:1. The joint section of the socket was redesigned based on the use of the internal adaptors.2.

The major changes from the design conceptmodel are as follows:1. The joint section of the socket was redesigned based on the use of the internal adaptors.2. The outer shape of the socket was modified at the bottom of the joint section, which nowstretches along the shape of the leaf spring.3. The internal structure has been improved to increase structural strength.4. The form of the outer trabecular structure was modified greatly.1762fig. 21 The design sketchRegarding the item 2 above, it was estimated that the stretched section would functionas a strength structure when the prosthetic leg is compressed downward. Simultaneously, thismodification aimed to enhance a sense of unity of the appearance of the prosthetic leg when itassembled. For item 4, the number of holes on the outer trabecular structure was reduced toreinforce the strength. Although the lightweight properties of the socket are significant, thedesign prioritizes reliance on strength at first.

Although the lightweight properties of the socket are significant, thedesign prioritizes reliance on strength at first. Fig. 22 and 23 indicate the results of CADmodeling and the appearance of the socket after the production.fig. 22 The CAD modelfig. 23 The manufactured socket1763The strength experiment for the running test modelA strength experiment for the running test model was conducted. The universal strengthtesting machine Autograph AG−IS (Shimadzu Inc.) was used for the test. Two pieces of jigs wereused in the test (fig. 24). One is made of aluminum and designed to hold the socket vertically onthe machine. The other is made by AM and its shape is designed to fit into the socket. The resultof the experiment is shown in fig. 25 underneath. It has been confirmed that the running testmodel can withstand a load of over 17kN (176,594N), satisfying the strength benchmark that wasoriginally targeted.fig. 24 The strength experimentfig.

24 The strength experimentfig. 25 Load vs stroke trace,X-axis refers stroke, Y-axis refers loadThe primal running testThe primal running test was conducted by using the running test model prostheticsocket. Scenes of the test are shown in fig. 26. The main purpose of this test was to get feedbackfrom the athlete. Although the structural strength of the socket itself was certified in a previouscompression test, the reliance of the locking mechanism that fixes the residual limb and thesocket together was not confirmed to be completely reliable. Therefore the maximum speed wasset at a jogging pace.

Therefore the maximum speed wasset at a jogging pace. The knowledge that we have obtained through the test is listed andcategorized as follows:1764Weight (fully assembled, include the leaf spring)Existing RSP 1449gAMed RSP running test model 1505gStructural issuesThere is no problem with either the strength or bending stability.No deformation around the aluminum adaptors was seen.The locking mechanism between residual limb and socket can be seen as lacking reliability.Friction of surfaceThe socket surface has a lot of friction; it may be uncomfortable while attaching theprosthesis.The end of the residual limb felt uncomfortable when attaching the socket.

It could becomepainful for the amputee.The center of gravity:Is behind the center of the socket and affects the user’s notion of speed, making the user feelas if they were running slower.AppearanceOrganic styling and characteristics.Coloring the external surface would be ideal for personalization purposes.Giving it an agile look would be preferable.fig. 26 The primal running test1765ConclusionA new design concept for RSPs was established. The concept enables to manufactureboth functional and aesthetically pleasing prosthetic socket by utilizing the benefits of AMtechnology. We succeeded in manufacturing a prosthetic socket that can be used for the primalrunning test by using a Nylon 12 based LS type AM machine.Remaining issuesOne of the characteristics of objects that are manufactured by AM is the unevenness ofthe surface. The surface has a lot of friction and it could become a negative factor for the user.

The surface has a lot of friction and it could become a negative factor for the user. Inorder to improve the product value, new methods should be established for an effective treatmentof the rough surface. It is also desirable that the method can cope with complex design forms.Regarding the locking mechanism between the socket and the amputee’s residual limb(fig. 26 left picture), the use of existing metal parts cannot be avoided. Investigation for jointmethods that are more supportive and durable than the current options should be required interms of size calibration and material strength.The purpose of this study is to provide mass-customized products for more peoplerather than providing tailor-made products for specific persons. As a next phase of this research,we have planned the development of a CAD software that is able to apply the method of designobtained by this work to prosthetic limbs.

signing Material Properties Locally with Additive Manufacturing technology SLM A.B. Spierings*, K. Wegener†, G. Levy*INSPIRE AG – institute for rapid product development irpd, Lerchenfeldstrasse 5, CH-9014 St.Gallen, Switzerland † ETH Zürich, Institut für Werkzeugmaschinen und Fertigungstechnik (IWF), Tannenstrasse 3, CH-8092 Zurich, Switzerland AbstractAdditive Manufacturing technologies are known to allow the production of parts with an extreme degree of complexity, enabling design and functional part optimization. So far the development of processing parameters and analyze of corresponding materials focuses on dense materials for maximized material properties. However, AM processes like Selective Laser Melting, allow also the generation of materials with some degree of porosity affecting their mechanical properties. A DOE was set up for SLM processed SS 17-4PH / AISI-630 material with porosity between 0% and about 26% in order to analyze mechanical properties.

A DOE was set up for SLM processed SS 17-4PH / AISI-630 material with porosity between 0% and about 26% in order to analyze mechanical properties. The results presented show that the porosity significantly affects material ductility and hardness, offering the possibility to design a material according to the required mechanical behavior of the parts produced. Therefore, this AM enabling features allows a multi-property component design by appropriate local parameter setting.Keywords: Additive Manufacturing, Selective Laser Melting, stainless steel, material propertiesI. IntroductionAdditive Manufacturing Technologies offer many advantages for the tool-less production of highly complex structured parts directly from CAD data. Especially the Selective Laser Melting (SLM) Technology, a powder-bed based process, allows the production of parts in a wide range of metallic materials by selectively fusing a metallic powder material.

Thereby, the remaining part-porosity can be kept below 1% by choosing appropriate processing parameters. The resulting static mechanical properties are typically widely comparable to values from conventionally processed materials, e.g. for stainless and hot-work steel [1, 2], Aluminium [3-5], Titanium [6], Ni-based materials [7, 8]. Consequently, such parts can be used in a wide field of industrial applications: Tooling industry, lightweight structures for automotive and aerospace, functionally optimized parts for diverse industrial and customized implants and instruments for medical applications. The SLM process uses a high power Nd-YAG laser source in order to continuously create side by side scan tracks filling the successive cross-sections of the parts to be produced. The main processing parameters are therefore the power of the laser source, the scan velocity, the hatch distance of the laser scan tracks and the thickness of the powder layers to be scanned [9].

The main processing parameters are therefore the power of the laser source, the scan velocity, the hatch distance of the laser scan tracks and the thickness of the powder layers to be scanned [9]. An almost densely processed material can be reached for the successful selection of a suitable set of these processing parameters. Typically, there is not only one possible set of processing parameters for a given material, but several different combinations. Design of Experiment (DOE) techniques are therefore useful to analyse the significance of the different processing parameters and to select suitable combinations for a given target value. Different researchers have used DOE in order to analyse the SLM-process and its parameters. Chatterjee [10] used a central composite design to analyse the effect of the main processing parameters on the part density / porosity and the hardness of a M2 work steel.

However, as AM Techniques are master forming technologies [16], not only the macroscopic geometrical structure but also the material itself with its properties are defined by the process. Therefore, by adjusting processing parameters, a desired degree of porosity can be generated affecting the mechanical properties of the material, e.g. its elasticity. The major effect of porosity on E-Modulus has been widely investigated and different models have been proposed [17-19]. Current work aims at investigating the possibilities for adjusting basic mechanical properties such as E-Modulus, hardness for a SLM manufactured stainless steel 17-4PH / AISI-630 material by using a full factorial DOE. The dependency of E-Modulus on porosity is validated against the model of Boccaccini [18]. The results show that adjustments of mechanical parameters are possible in a wide range, especially for hardened 17-4PH material.

The results show that adjustments of mechanical parameters are possible in a wide range, especially for hardened 17-4PH material. This offers the opportunity to locally design the mechanical behaviour of a structure not only by its geometrical design, but also by adjusting the material porosity.II. Methods and Materialsa) SLM Machine and MaterialsThe SLM machine type Concept Laser M2 was used, which is equipped with a Nd:YAG fiber laser having a maximal laser power of about 190W at the build platform and a nominal laser spot diameter of about 0.1mm. The scan strategy used to produce the tensile samples is a chess-board like scanning structure with 5x5 mm2-squares. More details are described in [20, 21]. A hardenable stainless steel 17-4PH / AISI-630 material was used for this study.

More details are described in [20, 21]. A hardenable stainless steel 17-4PH / AISI-630 material was used for this study. The raw powder material had the characteristics according to Table 1, allowing a productive and high quality production at all layer thicknesses used (see Table 2) [2].Table 1: Composition (%) and characteristics of the powder material used Composition Fe bal. Cr 16.7 Ni 4.5 Cu 4.3 Si 0.4 CbTa 0.3 Mn 0.3 N 0.12 C 0.02 PowderD10 = 16.4µmD50 = 26.8µmD90 = 42.7µmHalf of the specimens were heat treated according to the following procedure: Annealing at 1350°C, solution heat treatment at 1050°C, subsequent deep freezing & ageing at 480°C / 1h (condition H925).b) Design of ExperimentA full-factorial experimental design was chosen. The factors and corresponding levels are listed inTable 2.

The factors and corresponding levels are listed inTable 2. The levels were defined on the basis of the results of pre-trials.Table 2: DOE factors and corresponding levels( - ) Build orientation Layer thickness tLayer (µm) Scanning speed vScan (mm/s) Laser Power1 PLaser (W) (-) AnnealingLevel 1 Vertical (90°) 30 800 105 withoutLevel 2 Horizontal (0°) 50 1300 190 withThe Hatch distance was kept constant at 0.0975 mm (35% overlap of scan tracks). Cylindrical blanks for tensile test specimens (Figure 1) were produced with all combinations of above factors with 2 repetitions for each combination. Additionally, each combination was produced without and with heat treatment (annealing), resulting in total 96 specimens.

Additionally, each combination was produced without and with heat treatment (annealing), resulting in total 96 specimens. The specimens where then drilled to the final geometry according to ISO/DIN 50’125-B4x20.Figure 1: Tensile specimens (190W, 800mm/s) – horizontally produced blank (left) and final drilled specimen according to DIN-50125-form B (right).SFF 2012Review PaperTensile testing was performed using a universal testing machine type Zwick-1484 with a pre-load of 20MPa and a testing speed of 10-3 s-1. Brinell hardness measurements (HB 2,5/62,5) were performed at the end planes of the final tensile specimens. Therefore, for the horizontally built specimens, hardness was measured in parallel to the layers where as for the vertically build specimens, hardness was measured vertically to the layers.III.

Therefore, for the horizontally built specimens, hardness was measured in parallel to the layers where as for the vertically build specimens, hardness was measured vertically to the layers.III. Resultsa) Material density Density was measured on pre-manufactured cubes (10x10x10 mm3) using the Archimedes method, 3) supplied to the powder layer was calculated offering good repeatability [22]. The energy density (J/mm according to equation 1 [23], taking into account laser power PLaser (W), scan speed vscan (mm/s), hatch distance hs (mm) and powder layer thickness tLayer (mm): P Laser ⋅ ⋅ th sE−density=(1)vscanLayerFigure 2 shows the influence of the specific energy input into the powder layer on the resulting part density.

Good curve fitting for a resulting part density >≈ 75% was obtained with equation 2, where a = 1/E-Density (r2 = 0.96).Density=92.0+2.12−⋅ a514⋅a2+4318⋅a3(2)Figure 3 shows exemplary the pore distribution and sizes for porous (left) and almost dense material. It can be noted that with insufficient energy input, the pores become larger and partially interconnected.Figure 2: Part density versus Energy density (J/mm3). Confidence level 95%Figure 3: Micrographs, left: PLaser = 190W, vScan = 1300mm/s, tLayer = 50µm, right: PLaser = 190W, vScan = 800mm/s, tLayer = 30µm.E-modulus, hardness and Yield strengthLiterature values for the E-modulus for wrought 17-4PH material are between 197 – 207GPa (condition H925) [24]. The E-Modulus of the AM processed, heat treated 17-4PH material is highly comparable: For optimised density 199 GPa (horizontal) and 187 GPa (vertical), respectively.

The E-Modulus of the AM processed, heat treated 17-4PH material is highly comparable: For optimised density 199 GPa (horizontal) and 187 GPa (vertical), respectively. This difference between horizontal and vertical build orientation is about 5.7%; the absolute differences remain comparable over the whole density range (9.9 ± 2.9 GPa). Figure 4 shows that the modulus for heat treated material highly depends on the material density: Allowing a porosity of about 26%, E- modulus can be reduced to only ≈ 1/5 of the maximal value.SFF 2012Review PaperIn contrast, the modulus of the non-heat treated samples varies significantly around a mean value of ≈ 152 GPa and no significant trend in its relationship to porosity can be identified (Figure 4).

The stiffness values for the 74% samples could not be measured accurately and they broke very early.Figure 4: E-Modulus for heat treated and non-heat treated material versus density, vertical & horizontal build orientationHardness is also dependent on the material density for both heat treated and non-heat treated material (Figure 5). For dense, heat treated material, hardness values are within literature value ranges [24] and statistically not significantly different for horizontal or vertical build orientation, respectively (horizontal: HB 400 ± 1.6, vertical: HB 389 ± 16.3).

For the non-heat treated material the maximal values are also almost identical (horizontal: HB 233, vertical: HB 235).Figure 5: Hardness for heat treated and non-heat treated material versus density, vertical & horizontal build orientationIn addition to E-modulus and hardness, Yield strength shows a good correlation to density or applied energy density (see Figure 2), both in heat treated and non-heat treated condition. Differences between horizontal and vertical build orientations exist, but are not significant for the heat treated condition. However, in the as-built condition, differences can be observed and remain comparable over the whole analysed range (108 ± 41 MPa).

However, in the as-built condition, differences can be observed and remain comparable over the whole analysed range (108 ± 41 MPa). This is a direct result of the dendritic microstructure having a long grain axis parallel to the z-axis (90° orientation) and can be understood by the well known Hall-petch relationship.SFF 2012Review PaperFigure 6: Yield strength - Energy density for heat treated and non-heat treated condition, vertical and horizontal build orientationIV. DiscussionThe energy density (equation 1) can be considered as the main driving parameter for material density. Presented data (Figure 2) are valid for all combinations of laser power, scan velocity and layer thickness according to Table 2. Therefore, no distinction is made between specific factor combinations and observations in subsequent presented results.

Therefore, no distinction is made between specific factor combinations and observations in subsequent presented results. The following discussions are focusing on heat treated 17-4PH material, as the results for the E-modulus (Figure 4) show a very good correlation to material density. Nevertheless, the basic behaviour of not heat treated material can be compared to the heat treated material, except for the E-modulus (Figure 4). The dependency of E-modulus on material density, for both horizontal and vertical build orientation, can be described by a 2nd degree polynomial fit (equation 3, Figure 7).E-modulus = - 540952 + 730482 * density + 1037193 * (density-0.9371)2(3)Figure 7: E-modulus with 2nd degree polynomial fit (r2 = 0.99), 95% confidence interval for expected mean valuesHowever, it is interesting to compare this porosity-dependent behaviour of E-modulus to commonly used mathematical models; Boccaccini [18] gives a good overview on a variety of different models.

Typically, these models have been developed for other material classes (e.g. ceramics) and corresponding processing techniques. Most of the models assume spherical and homogeneously distributed pores of the same size. However, this is not true in most of the real cases. For AM processed materials, a more or less homogeneous pore distribution can be expected, especially at low porosity levels (Figure 3 right, [22]). At higher porosity levels, pores begin to be interconnected, whichSFF 2012Review Paperis a result of the non-equilibrium process characteristics, leading to balling phenomena [25] of the scan tracks and insufficient melting of powder particles etc.

Therefore, pores will not be of the same size nor spherical in shape (Figure 3, left).Anyway, Figure 8 shows good correlation between caluclated and measured E-modulus using the formula presented by Baccaccini (equation 4), with P the material porosity (= 1 – density), EM the E- modulus of the pore-free material and R the particle size ratio, used in this study as a curve fitting parameter, giving insight into a (theoretical) structure of the pores. The smaller the R-value, the smaller are the pores.

The smaller the R-value, the smaller are the pores. Thus for a given porosity and a lower R-value, more but smaller pores are present in the material.Figure 8: Caluclated (Bacciaccini [18]) versus measured E-modulus for horizontal () and vertical (--) build orientation with 95% confidence interval for expected mean values.EP=EM⋅1( ) P − P −+ 1(2⋅ R ) RP ⋅(4)For EM = 194.5 GPa (mean for horizontal and vertical build oritentation) and R = 0.21, a very good = ), indicating general usability of conventional correlation can be found ( mathematical formalism for E-modulus. Boccaccini used R-values of 0.6 to 1.0 for ceramic materials, indicating bigger pores compared to AM processed materials. Material hardness highly depends on density (Figure 5), both for heat treated and non-heat treated material. However, below a certain density (≈ 87%), hardness is not further reduced.

Material hardness highly depends on density (Figure 5), both for heat treated and non-heat treated material. However, below a certain density (≈ 87%), hardness is not further reduced. This indicates that hardness is influenced by the pore structure and therefore by the deformation behaviour of the remaining material under pressure. Supposing isotropic material behaviour under tension and compression, the stiffness of this remaining material is affected by E-modulus for elastic deformation, and Yield strength for the beginning of plastic deformation. It is therefore reasonable to analyse hardness in relation to E-modulus and Yield strength (Figure 9). Detailed results on the dependency of Yield strength on material density is being presented in an upcoming publication.2 r vertical2 r horizontal=,98.099.0SFF 2012Review PaperFigure 9: 95%-Density-Ellipse for Hardness (heat treated) in relation to E-modulus and Yield strength.

R-values indicate correlations (restricted maximum likelihood method)For heat-treated material, a correlation to both parameters is obvious. In contrast, as E-modulus for the non-heat treated material does not correlate with porosity (Figure 4), only a correlation to Yield strength can be observed (Figure 10).Figure 10: Hardness versus Yield strength (non-heat treated). 95% confidence interval, R2 = 0.80V. ConclusionsSelective Laser Melting allows not only the production of highly complex three dimensional parts. Its peculiarity of generating the material in the sense of a master forming technology allows also designing the mechanical material properties of the parts, e.g. the stiffness, by allowing a specific degree of porosity. The dependency of the E-modulus on porosity can be described by known mathematical formula, e.g. by Boccaccini [18]. It is expected that this formula is easily adaptable to other materials.

The dependency of the E-modulus on porosity can be described by known mathematical formula, e.g. by Boccaccini [18]. It is expected that this formula is easily adaptable to other materials. By a space-resolved adjusting of the relevant processing parameters, the results show that it is possible to locally design not only the structure but also the specific material behaviour. This allows optimizing the global mechanical part behaviour before the desired technical requirements of a part in service. It opens the way to functionally improved medical implants or instruments, and generally to a wide field of machine parts etc.SFF 2012Review PaperVI. OutlookIn this publication, only density-dependent results on the E-modulus and hardness are reported. Further results on the dependency of mechanical properties of material density, especially for Yield strength and Ultimate strength, are going to be published in a separate publication.

Further results on the dependency of mechanical properties of material density, especially for Yield strength and Ultimate strength, are going to be published in a separate publication. Beside an analysis of 17-4PH material, other materials should be analysed in order to be able to transfer the possibilities of designing the mechanical behaviour of a structure into industry specific applications: Titanium, Aluminium, etc. Complementary, the dependency of the dynamic mechanical properties on density should be analysed in more detail and correlated to existing formula for the dependency of fatigue behaviour on porosity.AcknowledgementsThe authors gratefully acknowledge Mr. A. Frauchiger for his support in the production of testspecimens. We furthermore thank Dr. M. Schmid for his critical reading this paper.Literature1.2.3.4.5.6.7.8.9.10.11.12.13.14.15.Averyanova, M. and P. Bertrand.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDETERMINATION AND COMPENSATION OF THE SHRINKAGE BEHAVIOR OF CYLINDRICAL ELEMENTS IN THE FDM PROCESST. Koers1,2*, B. Magyar1,21 Chair of Design and Drive Technology (KAt), Paderborn University, Germany 2 Direct Manufacturing Research Center (DMRC), Paderborn University, Germany * Corresponding author: thorsten.koers@upb.deAbstractFused Deposition Modeling (FDM) is an additive manufacturing process to produce complex thermoplastic geometries layer by layer. The filament is melted in a nozzle, iteratively deposited, and then cools down. Due to the solidification process, the deposited filament strands deviate from their intended position due to shrinkage, resulting in significant geometric deviations in the final part.

Due to the solidification process, the deposited filament strands deviate from their intended position due to shrinkage, resulting in significant geometric deviations in the final part. In terms of dimensional accuracy, there is a need for optimization, especially for local curved geometries in relation to the global part with higher nominal dimensions. The aim of this study is to investigate the size and shape deviations for cylindrical FDM elements and to compensate the expected deformations by using an in-house software with adaptive scaling factors in the x-y plane. Previous studies mainly focus on simple, non-curved objects, this study also considers the influence of curvature and global as well as local deviations on the final part.IntroductionAdditive manufacturing (AM) is characterized by its layer-by-layer buildup and requires no tooling.

Without the use of a tool, there are almost no restrictions on the design of a part, such as for complex geometries that could not be manufactured in conventional processes. The basis for each AM part is its CAD model, from which the layers are separated and then recombined during manufacturing.Fused Deposition Modeling (FDM) is a widely used plastics extrusion process in AM for the production of prototypes, tools and final products. In this process, the layers of a part are deposited one after the other by a melted thermoplastic filament strand. The continuous strand is pulled by motors into a nozzle head, where it is melted and then deposited in a defined position onto the underlying part. The newly deposited layer and the layer underneath bond by means of thermal fusion, cool down and strengthen. The layers are now permanently joined. The problem is that the deposited strands deviate from their defined deposition position because of solidification.

The layers are now permanently joined. The problem is that the deposited strands deviate from their defined deposition position because of solidification. Depending on the part geometry the shrinkage occurs in the x/y/z direction with significant inhomogeneity. This results in a significant influence on the dimensional and shape accuracy of the part, which is usually required by high and constant quality in the final application [1].These geometric deviations of FDM parts have often been investigated in studies. Mostly, for simplification purposes, non-complex standard elements are used, which can be divided into curved and non-curved [2]. Mainly, the studies of geometrical deviations in FDM involve non- curved parts such as cuboids. The results show that the outer length has a significant influence on the dimensional deviations. The higher the nominal dimension, the higher the deviation. For curved38elements, investigations on cylinders and bores can be found.

The higher the nominal dimension, the higher the deviation. For curved38elements, investigations on cylinders and bores can be found. The results regarding dimensional accuracy show that inner dimensions are undersized and outer dimensions are oversized. Analog to the non-curved elements, it is shown that the deviations increase with increasing nominal diameter [3-5]. A countermeasure is offered by scaling factors, which are applied to the CAD model by the supplier or the software in the pre-process to compensate the deviation from the nominal geometry in the manufacturing process. Here, too, increased dimensional deviations occur with increasing nominal length. The reason for this is a constant scaling factor, which is independent of the nominal dimension. Particularly for higher nominal dimensions above 50 mm, there is a need for optimization regarding the geometric accuracy. Also, no local scaling is possible to adapt for structures with complex geometries [6-10].

Also, no local scaling is possible to adapt for structures with complex geometries [6-10]. KNOOP [6] has developed scaling factors adapted to the expected deviation based on experimental investigations. The necessity of a simulative approach is pointed out since the experimental parameter development is expensive. In addition, the adapted scaling factors for linear elements are not applicable to curved elements, since here an analog scaling factor would further amplify existing undersize [6]. For cylindrical parts there have been investigations that developed a method to compensate the shrinkage of holes in FDM parts using adapted internal infill structures [11-12]. The combination of curved and non- curved parts has been neglected in studies to this point. Furthermore, the correlation between global and local deviations is missing. Using the scaling factors, only simple parts are considered globally. Complex parts with local variations are not considered.

Using the scaling factors, only simple parts are considered globally. Complex parts with local variations are not considered. But dimensional deviations occur especially in parts that are predestined for AM due to their complex geometry.The aim of this study is to determine and then compensate the shrinkage of cylindrical parts in the FDM process by applying optimized scaling factors in the x-y build direction. For this purpose, a self-developed software with adapted scaling factors is investigated on cylindrical test specimens and iteratively optimized. A nominal dimension-dependent scaling is set to generate locally instead of globally targeted adaptations for a complex part for validation.Experimental set-upThis chapter describes the experimental investigations carried out to examine the geometrical accuracy of simple curved FDM parts. Therefore, a Stratasys Fortus 400mc is used, which is available at the research center.

Therefore, a Stratasys Fortus 400mc is used, which is available at the research center. The machine enables processing of high-temperature plastics such as ULTEMTM 9085 resin due to its heated build chamber. The material under investigation is ABS-M30, which is chosen because it is an inexpensive standard material for the Stratasys machine used and a typical application plastic that serves for comparability with other studies. The process parameters used for the ABS-M30 material are listed in Table 1.

The measuring procedure is carried out with a Nikon Altera 8.7.6 coordinate measuring machine as well as an outside micrometer [16]. The coordinate measuring system is used for tactile measurement with a measuring head diameter of 2 mm and a layered measuring strategy (5 measuring rows, 10 mm intervals). The measuring points defined for the outside micrometer are the lowest, middle, and uppermost planes in the x- and y-directions (see Figure 1). The z-direction of the cylinders is not considered in this study. The testing plan includes the number 3 per build job.Figure 1: Measuring points in x- and y-direction of the outside micrometerFour different scalings of the cylinders are considered in this study. Unscaled test specimens with the scaling factor (SFx,y) equal to one are used as a basis. Standard scaled test specimens with SFx,y = 1.0059, which are defined by Stratasys, are used as a second reference.

Standard scaled test specimens with SFx,y = 1.0059, which are defined by Stratasys, are used as a second reference. Subsequently, an in-house developed software with self-adapted scaling factors is applied to unscaled cylinders in the pre-process before slicing. The method and functionality of this software were presented by HECKER at the Polcom 2022 in Bucharest [17]. It is based on the theory that raster lines shrink primarily in the direction of their nominal length. By adjusting the individual raster lines the software targets to compensate the shrinkage behavior in each layer. In this study, the scaling of the variants "Adaptive Scaling V2.4" and "Adaptive Scaling V3.2" are examined for one build job each. The main distinction between these two variants is that the iteration "Adaptive Scaling V3.2" is compensating the curvature of the specimens and "Adaptive Scaling V2.4" is not.

The main distinction between these two variants is that the iteration "Adaptive Scaling V3.2" is compensating the curvature of the specimens and "Adaptive Scaling V2.4" is not. Finally, the variants "unscaled" and "Adaptive Scaling V3.2" are compared with each other using a demonstrator in the form of a bracket (see Figure 2). The bracket has a complex geometry which can only be manufactured with AM. It contains 20 individual non-curved and curved shapes with various length dimensions and outer diameters. These individual geometries are measured to check the global as well as local geometric deviations.Figure 2: CAD-Model of the demonstrator40ResultsThe influence of the investigated scaling factors on the dimensional deviations as a function of the nominal diameter is analyzed in Figure 3 a). The dotted lines represent a trend line and the indicator bars the minimum and maximum deviations of the measured values.

The dotted lines represent a trend line and the indicator bars the minimum and maximum deviations of the measured values. For the unscaled test specimens, strong dimensional deviations can be seen. The continuously increasing undersize with increasing nominal dimension confirms the previously described hypotheses as well as the statements of KNOOP [6]. With increasing nominal size, an increasing scattering of the measured values can be seen, which are from the nominal diameter of 35 mm above the standard deviation. Considering the shape accuracy of the unscaled cylinders, the so-called clover effect can be observed (see Figure 5a). A shape deviation occurs in each position at 45° to the x- or y-axis. At this position, the deposited raster lines have the highest length and accordingly the highest tensile force. This results in the strongest shrinkage in the strand direction, which influences the layer contour.

This results in the strongest shrinkage in the strand direction, which influences the layer contour. The raster alignment alternates layer wise with 90°, so the deformation occurs rotation- symmetrically at four positions over the entire cylinder geometry. This shape determines the characterization of the clover. In comparison, the standard scaled test specimens exhibit an improvement in the measured values. A nearly constant negative dimensional deviation as the nominal diameter increases can be observed. Above the nominal dimension of 35 mm an increased measured value scatter above the standard deviation occurs.

Above the nominal dimension of 35 mm an increased measured value scatter above the standard deviation occurs. Compared to the unscaled test specimens there is an analog occurrence of the clover effect.Figure 3: Influence on the resulting dimensional deviations of cylinders due to a) the different scaling factors including all measuring points (left diagram); b) the measuring height in z-direction for scaling factor “Adaptive Scaling V3.2” (right diagram)The graph of the measured values "Adaptive Scaling V2.4" shows the results of the test specimens with the in-house developed scaling factors. Here, continuously increasing negative dimensional deviations are obtained with increasing nominal diameter. Compared to the unscaled test specimens, a reduction of the deviations can be achieved. At 15 mm diameter, the difference is 0.10 mm and increases continuously up to 0.16 mm at 65 mm. The dimensional accuracy of the standard scaling cannot be achieved.

At 15 mm diameter, the difference is 0.10 mm and increases continuously up to 0.16 mm at 65 mm. The dimensional accuracy of the standard scaling cannot be achieved. Therefore, newly adapted scaling factors “Adaptive Scaling V3.2” are elaborated by HECKER in an iteration process. The iteration achieves a nearly41analog result to the standard scaling for the nominal diameter 15 mm. The dimensional deviation then increases with increasing nominal diameter up to 0.03 mm at 65 mm. Measurement value outliers are present to an increased extent. The cause of the inconsistent measured values within the standard deviation can be seen in Figure 3b. There, the dimensional deviations from “Adaptive Scaling V3.2” are shown as a function of the nominal diameter for the three measuring points top, middle, and bottom. It can be seen here that the test specimens have a conical curvature along the z-direction.

It can be seen here that the test specimens have a conical curvature along the z-direction. The deviation of the diameter at the lower section of the cylinder is almost the same as the values for the standard scaling. The measured values increase for nominal dimensions above 25 mm for the measuring plane in the middle as well as at the top. In contrast, an enhancement of the clover effect is visible in particular for the high nominal dimensions (see Figure 5b).Figure 4: Influence on the resulting a) roundness deviations (left diagram) and b) cylindricity deviations (right diagram) of cylinders due to the different scaling factorsIn terms of shape accuracy, almost parallel results can be seen for the scaling variants "Unscaled", "Standard scaling" and "Adaptive Scaling V2.4". "Adaptive Scaling V3.2" shows contrasting measured values. Figure 4 presents these graphically with a) for roundness deviation and b) for cylindricity.

"Adaptive Scaling V3.2" shows contrasting measured values. Figure 4 presents these graphically with a) for roundness deviation and b) for cylindricity. The unscaled test specimens show an increasing roundness deviation of 0.1 mm to 0.14 mm with increasing nominal diameter. The values for "Standard scaling" and "Adaptive scaling V2.4" run parallel with a difference of an additional 0.5 mm respectively 0.7 mm deviation. "Adaptive Scaling V3.2" has a high roundness deviation of 0.29 mm for the smallest diameter and decreases to 0.19 mm for diameters up to 65 mm (see Figure 4a). The same is observed for the cylindricity deviation. "Unscaled", "Standard scaling" and "Adaptive scaling V2.4" show parallel deviation trends, while "Adaptive scaling V3.2" shows a high deviation for the smallest diameter and decreases with increasing nominal diameter (see Figure 4b).

The high measured value scatter for all scaling variants is remarkable for the shape accuracy.42Figure 5: Visible roundness deviation in x-y direction for a) “Standard scaling”, d = 15 cm (left) and b) "Adaptive Scaling V3.2", d = 35 cm (right)The factors "Adaptive Scaling V3.2" applied to the demonstrator bracket for validation achieve the expected dimensional deviations for the existing curved geometries according to the preliminary investigations. An overall low measured value scatter is also identifiable, which leads to constant dimensional values. This is also shown in particular by the measured length dimensions for non-curved geometries of the bracket in both the x- and y-directions. Local as well as global dimensions were considered in the non-curved measurement. In comparison, the measured dimensional deviations of the unscaled brackets increase with increasing nominal dimension from local to global.

In comparison, the measured dimensional deviations of the unscaled brackets increase with increasing nominal dimension from local to global. The average value of the dimensional deviations across the entire part is -0.43 mm for the unscaled demonstrator (standard deviation: 0.25 mm) and -0.37 mm for the adaptive scaled demonstrator V3.2 (standard deviation: 0.18 mm). Considering the standard deviation, there is no significant improvement in the dimensional accuracy. It must be noted that both non-curved and curved geometries are included in these average values over different measurement heights in the z-direction. While "Adaptive Scaling V3.2" achieves constant measured values along the entire z-height for non-curved geometries [17], this accuracy is not achieved for curved geometries, as previously investigated using the cylindrical test specimens.

This can be seen in the measured value scatter of "Adaptive Scaling V3.2" on the curved geometries of the demonstrator bracket.ConclusionThe aim of this study was to investigate the size and shape deviations for cylindrical FDM elements and to compensate the expected deformations with optimized scaling factors in the x-y plane. For this purpose, cylindrical test specimens were investigated with adapted scaling factors of an in-house developed software. It has been shown that the considered approach to compensate the shrinkage per layer in the direction of the raster lines provides improved dimensional accuracy compared to unscaled specimens. The dimensional deviations for standard scaled test specimens were not achieved but were approximated by an iteration process. The adapted scaling factors achieve constant dimensional values in the strand direction. In the z-direction the dimensional deviations are locally low.

The adapted scaling factors achieve constant dimensional values in the strand direction. In the z-direction the dimensional deviations are locally low. Due to the formation of a conical shape of the cylindrical surfaces an increased measured value scatter results globally. A further iteration of the scaling factors is necessary for a constant dimensional value over the complete z-height. Furthermore, the circle approximation has to be modified so that the clover effect of the contour can be reduced. The validation on the bracket confirmed the results on the simple cylindrical test specimen and also for a complex part with varied geometry characteristics.43The presented approach enables an improvement of quality and consistency of dimensional accuracy in the FDM process. In further investigations, it is necessary to optimize the software so that the dimensional deviations are further reduced.

In further investigations, it is necessary to optimize the software so that the dimensional deviations are further reduced. The range of deviations should remain constant in order to further guarantee constant dimensional accuracy. In addition, the possibility of transferring the shrinkage behavior between different materials on the basis of specific material parameters should be examined in further analyses. This would significantly reduce the future experimental effort for modeling the shrinkage behavior of further materials. These open issues can be addressed in further investigations.AcknowledgementsThe presented results were obtained in the context of the research project ZI 723/12-1 “Simulation of the shrinkage behavior in Fused Deposition Modeling” sponsored by the German Research Society (DFG). We would like to thank the DFG for its support.References[1]A.

TERMINATION AND IMPROVEMENT OF BUILDING SPEED IN RAPID FREEZE PROTOTYPINGWei Zhang, Guanghua Sui, Ming C. LeuDepartment of Mechanical and Aerospace Engineering and Engineering Mechanics University of Missouri – Rolla, Rolla, MO 65409-0050 E-mail: mleu@umr.eduABSTRACTRapid freeze prototyping (RFP) is a solid freeform fabrication process that builds an ice part by rapidly freezing water in a layer by layer manner. One advantage of this process is the ability to build ice parts faster than other SFF processes. The factors that affect the speed of contour building and interior filling in RFP are identified. The influence of these factors is analyzed through heat transfer and material flow analyses. A model based on heat transfer analysis is proposed to determine the maximum achievable speed of contour building under stable conditions.

A model based on heat transfer analysis is proposed to determine the maximum achievable speed of contour building under stable conditions. Experiments are conducted to validate the performance of the proposed model for determination of building speed.KEYWORDS Solid Freeform Fabrication, Layered Manufacturing, Rapid Freeze Prototyping, Ice Patterns1. INTRODUCTIONRapid Freeze Prototyping (RFP) is a novel solid freeform fabrication (SFF) process that uses water as the building material and makes three-dimensional ice parts by freezing water according to their CAD models layer by layer. This process is being developed by us to address many of the existing issues in the current SFF processes, such as high cost, poor surface finish, low building speed, etc. Details of this process can be found in Ref. [1-4].

Details of this process can be found in Ref. [1-4]. The advantages of this process, including cleanness, lower cost, potential better surface finish and building accuracy, ease of stress compensation, and faster building speed, were discussed in [2-4]. Potential applications of this process include part design visualization and rapid tooling such as RTV silicone molding, investment casting, and sand casting[5-6]. The most promising application appears to be investment casting. In a patented investment casting process, called the Freeze Cast Process (FCP), ice patterns instead of traditional wax patterns were used to make metal parts. This process was developed to address problems and technical difficulties of traditional investment casting such as wax pattern expanding, ceramic shell cracking, etc.[5]. One major concern in the FCP process is ice pattern making. The traditional method of making ice patterns is by injecting water in a mold and making it frozen.

One major concern in the FCP process is ice pattern making. The traditional method of making ice patterns is by injecting water in a mold and making it frozen. Some main issues in this method include compensation of water expansion during freezing, air bubble removal, ice pattern de-molding, and part complexity limitations. With RFP it is possible to make accurate ice patterns directly from CAD models in a short time, without the cost and other issues of mold making.

With RFP it is possible to make accurate ice patterns directly from CAD models in a short time, without the cost and other issues of mold making. This is especially valuable in case that a small amount of complex metal parts and thus ice patterns are needed.RFP has the potential to increase the part building speed to a much higher level than mostother SFF processes because of the following reasons: 1) Ice layer binding is achieved naturally through hydrogen bonds, thus the environmental514temperature can be very low without the need to ensure sufficient layer binding strength through improving environmental temperature as other thermal deposition processes.2) Since water expands rather than shrinks when freezing, the expansion can be compensated by adjusting line spacing. As a result, there is much less stress caused by phase change as compared with other thermal deposition processes.3) Water has a lower viscosity than molten wax and plastic.

As a result, there is much less stress caused by phase change as compared with other thermal deposition processes.3) Water has a lower viscosity than molten wax and plastic. In conjunction with the above two features, RFP can use a faster interior filling mode by simply flushing water to the interior of a boundary (built with a regular scanning mode) to a predetermined height. The flushing mode is more efficient than the regular line-by-line filling mode. However, to take advantages of the above, the process parameters must be controlled to avoid building process failures such as the water line stability problem. To do this, a good understanding about heat transfer in building a new layer and the flow of deposited material is essential.

To do this, a good understanding about heat transfer in building a new layer and the flow of deposited material is essential. Through detailed thermal analysis and material flow analysis, parameters can be optimized to result in faster building speeds and to avoid failures in both contour building and interior filling in the meantime.The paper is organized as follows: Section 2 provides detailed heat transfer analysis and discusses the influence of various process parameters; Section 3 analyzes the material flow behavior and stability problem; and Section 4 concludes this paper.2. HEAT TRANSFER ANALYSISFigure 1 shows our current RFP experimental system. It includes: a) a three- dimensional positioning subsystem, b) a material extruding and ejecting subsystem, and c) a freezer. The whole experimental system is controlled by a personal computer, with the movements in the three axes and the nozzle ejecting frequency controlled by a PC- bus based stepping motor controller.

The whole experimental system is controlled by a personal computer, with the movements in the three axes and the nozzle ejecting frequency controlled by a PC- bus based stepping motor controller. The also material controlled by this motion controller. The material supplying subsystem feeds water to the ejecting nozzle from a reservoir by a pressure device at a pre-set constant pressure. The nozzle, which mainly consists of a miniature solenoid valve, a jet, and a heating cover, generates water droplets at a controlled frequency and size.extrudingsubsystemisFigure 1. Rapid Freeze Prototyping Experimental SystemThe formation of a well shaped and uniformly deposited ice line is fundamental to the building of a three-dimensional ice part. There are two dominating factors in the ice line building process: material freezing and spreading on the building surface. Freezing is caused by the fast cooling of the built layers and the cold environment.

Freezing is caused by the fast cooling of the built layers and the cold environment. It is a contributing factor to the improvement of ice line resolution. However, the spreading of the newly deposited water caused by surface interaction between the water and the top ice layer (associating with surface tension, capillary force, and viscous force) negatively affects ice line resolution. The balance between the positive factor and the negative factor determines the final ice line quality, i.e. line resolution and evenness. It is important to identify the parameters that have significant influence on the balance515and to understand this influence in order to control the process and build successful ice parts.

It is important to identify the parameters that have significant influence on the balance515and to understand this influence in order to control the process and build successful ice parts. Theoretical analysis of the heat transfer between a newly deposited layer and the previously built layers can provide quantitative information about the freezing process.To analyze heat transfer in both the newly deposited layer and the previously built layers, we need to treat boundary building and interior filling separately. Boundaries should be built with smaller layer thickness in order to get fine resolution, accuracy, and surface finish. Interiors can be built by flushing water every N layers (much thicker than one single layer) in order to improve the building speed. The newly deposited material for boundary building is mainly cooled down through conduction by the previously built layers, while convection is relatively insignificant.

The newly deposited material for boundary building is mainly cooled down through conduction by the previously built layers, while convection is relatively insignificant. For interior filling, in order to improve cooling rate, a cold metal plate can be placed on top of the newly filled layer. So the newly filled interior layer is cooled down mainly through conduction by both the metal plate and the previously built layers, while convection is again relatively insignificant. To simplify the analysis, we will not consider the heat transfer along the nozzle moving direction (i.e. scanning direction) and the heat transfer through convection. Since the boundary is built up with only one line, the line has no neighboring lines and thus no heat transfer in XY plane. For interior filling, the temperature gradient in XY plane is also negligible. So for both boundary building and interior building, heat transfer occurs mainly in Z direction and thus it is a one-dimensional problem.

So for both boundary building and interior building, heat transfer occurs mainly in Z direction and thus it is a one-dimensional problem. The analysis models are given in Figure 2(a) and Figure 2(b).Moving directionNew layerNozzleBuilt layersMetal plateSimplified cross sectionNew fillingMetal plateTop plate (only for interior building)Elements0 2HBuilt layersn(a) Boundary building model (b) Interior filling model (c) Finite elements Figure 2. Finite element analysis models for both boundary and interior buildingFinite element analysis is used to tackle this heat transfer problem. The model consisting of a multitude of previously built layers and a new layer is divided into small elements of three kinds (see Figure 2(c)): (A) interior elements (each with two neighboring elements), (B) surface element (on top of new layer), and (C) boundary elements (bottom and top elements touching the metal plate).

The governing equations can be obtained as follows[7]:T i(A) For an interior element P(i) (0<i<n) = 21(D+ 1+ )TF ( i 0T iTF ) i 0+ 1whereF 0=c(z2)(B) For the surface element P(0) in boundary building without a metal plate on top =D T TF )21( 0 0 0 (C) For the boundary elements P(0) and P(n)+ TF 102D…= D T T M 0 In the above equations,T n(with metal plate on top for interior building)D, TM, z, c, and lrepresent the updated temperature after time , metal plate temperature, element size in Z direction, specific heat, and conductivity ofiT5161iice/water, respectively.Figure 3 shows some heat transfer analysis results in boundary building. In this analysis, the time required to freeze the deposited material completely, called the freezing time and denoted by F, is investigated. Before F, at least part of the material is in liquid state and there are material flow and surface tension effects.

Before F, at least part of the material is in liquid state and there are material flow and surface tension effects. After F, heat transfer only lowers the new layer’s temperature, with no surface tension effect, no phase change, and no significant volume change. Figure 3(a) shows the ice percentage change in the new layer as a function of time. The freezing time is the time (0.62 sec.) when the freezing percentage reaches 100%. Figure 3(b) shows the cross-sectional temperature distribution at the freezing time.

layer thickness0.80.60.40.2e m T g n i z e e r Fi) c e s (0024681012604020Initial TemperatureEnvironmental Temperature(e) Freezing time vs. initial droplet temperature ((cid:176) C) (f) Freezing time vs. environmental temperature ((cid:176) C)Figure 3. Analysis results for boundary buildingWe can see that layer thickness and environmental temperature are more significant in affecting the freezing time, while initial droplet temperature has limited influence. This can be explained by the fact that water releases much more heat during freezing than during temperature drop from the initial droplet temperature to water’s freezing point. The height of built layers only51703 e m T 2 g n 1 i z e e r 0 Fi0.3) c e s (significantly affects the freezing time before it reaches about 2 mm. After that, its influence becomes negligible.

After that, its influence becomes negligible. This is good for process control because we do not have to adjust other process parameters with built height once it reaches 2 mm. The information obtained from the heat transfer analysis provides quantitative understanding of the temperature and phase change in depositing the new layer. It can help optimize process parameters and solve the stability problem discussed in the next section.Figure 4 shows the analysis result for interior filling. In this analysis, the time needed to cool down the new filling layer is investigated. This time, called the cooling time, is defined as the time required to cool down the new layer’s temperature to within 20% of the initial temperature difference with the metal plate’s temperature. For example, if the initial droplet temperature is 5(cid:176) C and the metal plate’s temperature is –20(cid:176) C, then the cooling time is the time that lowers the new layer’s highest temperature to –15(cid:176) C.

The reason that we use Cooling Time instead of Freezing Time for interior filling is that we must cool down the newly filled layer in order to start the next layer. It can be seen form Figure 4 that a smaller filling thickness or a lower metal plate temperature results in a shorter cooling time. However, considering the time required to move the metal plate over the filling layer, smaller layer thickness means more operations and thus not necessarily reduces the total filling time. The optimal filling thickness needs to be calculated with the consideration of both cooling time and metal plate operation time.) c e s (306e m T g n i l o o Ci201000123453525420 -15Layer Thickness (mm)Environmental Temperature(a) Cooling time vs. layer thickness (at -20(cid:176) C) (b) Cooling time vs. environmental temperature ((cid:176) C)Figure 4. Analysis results for interior building3.

layer thickness (at -20(cid:176) C) (b) Cooling time vs. environmental temperature ((cid:176) C)Figure 4. Analysis results for interior building3. MATERIAL FLOW ANALYSISAfter a droplet is ejected onto the building surface, it will deform, spread, and freeze. The shape change ends with complete freezing. Deformation and spreading help to make continuous ice lines, however, overspreading will cause failure of shape control and ice line building. The shape deformation and material flow in the plane perpendicular to the scanning direction (nozzle moving direction) is analyzed in Ref. [8]. The material flow along the scanning direction is discussed in this section.In our experiments of ice line and part building, we have observed the stability problem as shown in Figure 5. The left one generated a good ice line with even line width, while the right one lost control of line building (lots of water blobs were formed). This is a typical stability problem.

The left one generated a good ice line with even line width, while the right one lost control of line building (lots of water blobs were formed). This is a typical stability problem. The reason for this phenomenon is the complex interaction of water surface tension, capillary force, gravity, and water viscosity[9]. During the process that an evenly distributed water line changes to a spherical blob, the energies associated with this process are (refer toEL=E=ELkL Figure 6):R,+EE+R+EEs ¢=2 3Lkg,Eg=Lk g1 3Es=Lk s,E=Lufk)(518e m T g n i z e e r Fi) c e s (Therefore:= uf )(1 k[kks¢Lk1 3Lk g]32In the above,EL,ER,EEE,,Sgare the surface energy of the water line, surface energy ofthe water sphere, energy due to water viscosity, energy due to capillary force, and energy due to gravity, respectively. are the associated coefficients.

are the associated coefficients. These equations relatek,k,k,k,k¢gsthese various energies to the water line length L. For example, the energy consumed due to gravity is proportional to the sphere radius, and is thus proportional to the cubic root of the sphere volume or the water line length (since the water line width and line height are fixed).The above equations are based on the following assumptions:§ L >> R, and thus the line width change during the blob formation is negligible. § The initial and final shapes of water deposit are an ideal even line and a half sphere. § During the shape changing, there is no temperature, phase, or material property change. §)(ufis a monotonous increasing function about the average shape changing speed uRLFigure 5. Built ice lines (left: fine line; right: unstable line)Figure 6.

§)(ufis a monotonous increasing function about the average shape changing speed uRLFigure 5. Built ice lines (left: fine line; right: unstable line)Figure 6. Shape change analysis model++>+and u will be positive, which means the )(uf shape will change at a certain average speed u . Otherwise, the shape will stay stable or semi- stable. A greater u means faster shape change, and the water line is more unstable. It can be seen that when the water line length L increases, the corresponding increases of the five energies are not the same. Some energies increase proportionally and others increase relatively slowly. As a and u increase and the water line becomes more unstable. This has been result, both confirmed by our experimental observations.EEE, thenEEIfgLSR)(ufIn rapid freeze prototyping, since water freezes after deposition, the line profile is changing with time.

Accordingly, the percentage of water after deposition is a function of time and thus is a function of distance from the nozzle along the scanning direction. Figure 7 shows the water-ice boundary profile for a newly deposited layer along the scanning direction and the influence of layer thickness and environmental temperature on the boundary profile. Smaller layer thickness (Figure 7(a)) and lower environmental temperature (Figure 7(b)) result in a smaller amount of water in the deposited water line.zh(x)Scanning directionwatericexF = v Fxicewatericewater(a) = 0.1 mm, T = -20 C (b) = 0.1 mm, T = -30 C (c) = 0.15 mm, T = -20 C Water volume V(xF) = 1.47 mm3 V(xF) = 0.53 mm3 V(xF) = 6.4 mm3Figure 7.

Influence of layer thickness and environmental temperature on boundary profile and water volume519Therefor, with consideration of water-ice profile change (using volume instead of using length), as an analogy to the stability of a pure water line, the deposited water becomes more unstable when the amount of water becomes larger. Since the deposited water will eventually freeze, this translates to a larger height variation in the built part for a greater amount of water. The amount of water can be computed using the following equation: V(xF ) = w where v, w,x F[ (] - h(v )(cid:242)(cid:242)d = V(v F ) (1) 0 , F, xF, and h(x) are scanning speed, line width, layer thickness, freezing time, scanning distance at freezing time, and water percentage function, respectively. The function h(x) can be obtained by transforming the relationship of freezing percentage vs. time (as shown in Figure 3(a)) to freezing percentage vs.

The function h(x) can be obtained by transforming the relationship of freezing percentage vs. time (as shown in Figure 3(a)) to freezing percentage vs. scanning distance (x = v ).h(x)]dx =F[vw0To build an ice part, we need to control the height variation within an acceptable limit. To achieve the variation control, according to the above conclusion, we need to restrict the amount of water (V(xF)) in the deposited line within a limit. So for any given environmental temperature (i.e. given h(x)), there exists a maximum alloable scanning speed.

So for any given environmental temperature (i.e. given h(x)), there exists a maximum alloable scanning speed. A scanning speed larger than this speed limit will proportionally cause the amount of water (V(xF)) in the deposited line exceeding the limit (based on Equation (1)), and thus causing the variation over the acceptable limit.In order to quantitatively determine the maximum allowable scanning speed to obtain acceptable part height variation for any given environmental temperature, we need to experimentally determine a maximum scanning speed for one environmetnal temperature. The speed value will be used to compute the maximum allowed water amount V(xF). Once V(xF ) is obtained, we can use equation (1) and the freezing percentage function obtained from heat transfer analysis to calculate the maximum scanning speed for other various environmental temperatures.As an example, we consider building a cylinder with a diameter of 50 mm and a height of 100 layers.

The acceptable part height variation is 2%. The maximum scanning speed for the environmental temperature of –20(cid:176) C is experimentally determined to be 75 mm/sec., and the computed V(xF ) equals to 1.47 mm3. The maximum scanning speeds at other environmental temperatures are computed as shown in Figure 8, which also provides the experimentally measured maximum scanning speed at various environmental temperatures. It can be seen that the predictions agree well with the experimentally measured results. Thus the value of the fundamental understanding and analysis as described is allowing us to theoretically determine the maximum allowable scanning speeds at various environmental temperatuers by measuring the maximum scanning speed at only one temperature.Allowed Scanning Speed (mm/sec)12090Measured Predicted6030030 -20 -10 0 Environmental Temperature (°C)Figure 8. Predicted allowable scanning speed vs. experimentally measured allowable scanning speed5205.

Predicted allowable scanning speed vs. experimentally measured allowable scanning speed5205. CONCLUSIONThe heat transfer analysis together with material flow analysis helps to understand the freezing process and provides useful information to the selection of building parameters for build time reduction in rapid freeze prototyping. The analysis of ice line building stability provides an analytical method to predict maximum allowable scanning speed for any given environmental temperature with only one experimentally measured result, and this method is verified experimentally.Further analyses and observations (especially microscopic material flow and freezing) are needed in the future to provide better understandings about rapid freeze prototyping, and to improve building speed as well as building accuracy and surface finish.REFERENCES 1) Leu, M. C.

C. and Zhang, W., “Research and Development in Rapid Prototyping and Tooling in the United States,” Proceedings of the International Conference on Rapid Prototyping and Manufacturing, Tsinghua University, Beijing, China, July 21-23, 1998.2) Zhang, W., Leu, M. C., Ji, Z., and Yan, Y. N., “Rapid Freezing Prototyping with Water,” Proceedings of the 9th Annual Solid Freeform Fabrication Symposium, University of Texas at Austin, Austin, TX, pp. 185-192, August 1998.3) Zhang, W., Leu, M. C., Ji, Z., and Yan, Y. N., “Rapid Freezing Prototyping with Water,” Materials and Design, Vol. 20, pp. 139-145, June 1999.4) Zhang, W., Leu, M. C., Sui, G. H., and Ji, Z., “An Experimental and Analytical Study of Ice Part Fabrication with Rapid Freeze Prototyping,” Proceedings of the 10th Annual Solid Freeform Fabrication Symposium, University of Texas at Austin, Austin, TX, pp. 591-598, August 1999.5) Yodice, A., “Freeze Process Cuts Casting Costs,” Advanced Materials & Processes, Vol. 155, No.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDETERMINATION OF A SHAPE AND SIZE INDEPENDENT MATERIAL MODULUS FOR HONEYCOMB STRUCTURES IN ADDITIVE MANUFACTURINGThao Le*, Dhruv Bhate†, John M. Parsey, Jr.*, Keng H. Hsu*Arizona State University, Mesa, AZ †Phoenix Analysis & Design Technologies Inc. (PADT), Tempe, AZAbstractMost prior work on modeling cellular structures either assumes a continuum model or homogenizes “effective” cell behavior. The challenge with the former is that bulk properties do not always represent behavior at the scale of the cellular member, while homogenization results in models that are shape specific and offer little insight into practical design matters like transitions between shapes, partial cells or skin junction effects.

This paper demonstrates the strong dependence of measured properties on the size of the honeycomb specimen used for experimental purposes and develops a methodology to extract a material modulus in the presence of this dependence for three different honeycomb shapes. The results in this paper show that the extracted modulus for each shape converges as the number of cells in the specimen increases and further, that the converging values of the material moduli derived from the three shapes are within 10% of each other.Introduction: The Role of ModelingIt is well appreciated that cellular structures such as honeycombs and lattices are a particularly exciting research frontier in Additive Manufacturing (AM). Broadly speaking the research landscape in this space involves design, manufacturing, modeling and implementation into an application. Several software solutions enable design with cellular structures and the published literature on manufacturing them is rapidly growing.

Several software solutions enable design with cellular structures and the published literature on manufacturing them is rapidly growing. This work focuses on the modeling aspects of cellular structures produced with AM. Modeling in this context is the analytical representation of material behavior, primarily for use in predictive analysis. This is a critical aspect of enabling true simulation-driven design, where the design is the outcome of some objective such as stiffness- to-weight maximization relative to an allowable stress. From a more abstract point of view, modeling links up design and manufacturing capabilities to the application. Without models that describe cellular structure behavior, we have design tools that make structures that we can even manufacture, but we have little confidence in their ability to perform the desired function in the end-application, especially when these are critical to function.

While it is possible to certify components with cellular structures through rigorous testing, valid models and robust simulation together can both drive down qualification costs and unlock untapped potential that a normal design-manufacture-test cycle fails to uncover.Approaches to develop predictive models for cellular materials fall into three broad categories depending on the level of discretization at which the property is modeled: at the level of the individual unit cell, at the level of the connecting member or finally, at the level of each material2148point. The most straightforward approach is to use bulk material properties to represent what is happening to the material at the cellular level [1-4]. This approach does away with the need for any cellular level characterization and in so doing, does not have to account for size or contact effects described previously that are artifacts of having to characterize behavior at the cellular level.

However, the assumption that the connecting members in a cellular structure behave the same way the bulk material does can be particularly erroneous for AM processes that often introduce significant size specific behavior and large anisotropy. It is important to keep in mind that factors that may not be significant at a bulk level (such as surface roughness, local microstructure or dimensional tolerances) can be very significant when the connecting member is several millimeters thick or less, as is often the case. The level of error introduced by a continuum assumption is likely to vary by process: polymeric processes like Fused Deposition Modeling (FDM) are strongly anisotropic with highly geometry-specific meso-structures and an assumption like this will generate large errors.

On the other hand, it is possible that better results may be had for powder based fusion processes used for metal alloys, especially when the connecting members are large enough and the key property being solved for is mechanical stiffness (as opposed to fracture toughness or fatigue life).A common approach in the literature that attempts to account for behavior at these smaller scales is the use of homogenization, which involves representing the effective property of the cellular structure without explicitly accounting for the cellular geometry itself. This approach has significantly lower computational expense associated with its implementation in simulation software.

This approach has significantly lower computational expense associated with its implementation in simulation software. Additionally, it is relatively straightforward to develop a model by fitting a power law to experimental data [5-9] as shown in the equation below, relating the effective modulus E* to the bulk material property Es and their respective densities (ρ and ρs), by solving for the constants C and n.(1)𝑛𝑛𝜌𝜌 𝜌𝜌𝑠𝑠 �∗= 𝐶𝐶 𝐸𝐸𝑠𝑠 �𝐸𝐸While a homogenization approach is useful in generating comparative, qualitative data, it has some difficulties as a reliable material model in analysis and simulation. This is primarily because homogenization of cells only works for the specific cell in question - so every new cell type needs to be re-characterized.

This is primarily because homogenization of cells only works for the specific cell in question - so every new cell type needs to be re-characterized. Additionally, the homogenization of these cells obscures insight into how structures behave in the transition region between different volume fractions or at junctions of cellular members and walls, even if each cell type is calibrated at a range of volume.The third approach involves describing behavior not at each material point or at the level of the cell, but at a level in-between: the connecting member (also referred to as strut or beam) as shown in Figure 1. This approach has been used by researchers [10-12] and in our previous work as well [13] by invoking beam theory to first describe what is happening at the level of the member and then using that information to build up to the level of the cells.

This approach, while promising, has some challenges as well: from a computational standpoint, the validation of these models typically requires a modeling of the full cellular geometry, which can be prohibitively expensive. Additionally, the theory involved in representing member level detail is more complex, makes assumptions of its own (e.g. modeling the "fixed" ends of the beam). It is not clear at the present moment if this is justified by a significant improvement in the model's predictability compared to the above two approaches. This approach does however have the potential to be truly shape2149independent since it does not require the assumption of a specific unit cell design and is why it is the method of choice for this work as well.Figure 1.

Representation of cellular structures as an assemblage of connecting members such as the beam shown here enables the use of beam theory in the development of models (P and C represent loads, M the moment, θ and δ the angle and displacement of the beam)Modeling in the Presence of Shape and Size DependenceThere are several challenges associated with modeling cellular structures and materials. Chief among these is the fact that behavior of these structures is highly dependent on the shape of the specific cells used in the structure being studied [13] – in other words, a hexagonal honeycomb structure will behave differently than a square or triangular structure. What is less intuitive is that data obtained from test specimens used to calibrate models is highly dependent on the size of the specimen tested as well.

What is less intuitive is that data obtained from test specimens used to calibrate models is highly dependent on the size of the specimen tested as well. Consider the experimental data shown in Figure 2 for honeycombs tested in compression, showing that the predicted effective modulus increases with increasing number of cells in the axial direction, but reduces (at a lower rate) for increasing number of cells in the transverse direction. Other experimental and numerical (simulation) studies concur with this observation [14-19] and have even shown that the trend and sensitivity of modulus to the number of cells is itself dependent on the shape being studied [20].Figure 2. Effective modulus under compression showing a strong dependence on the number of cells in the structure.

Effective modulus under compression showing a strong dependence on the number of cells in the structure. (The image on the bottom right corner shows a honeycomb sample under a compression test with the Instron 8801 mechanical testing system).2150This work focuses on developing a method to establish a shape- and size independent modulus, building on previous work [12] that proposed a method for extracting a material modulus by applying established analytical theory [13] to measure effective behavior of hexagonal honeycomb structures. In this paper, the approach is extended to account for size of specimens (number of cells) and to validate shape insensitivity by comparing results obtained from hexagonal, triangular and square shapes.

While the work done here is using the Fused Deposition Modeling (FDM) process, the methodology is expected to be applicable for other manufacturing processes as well.Specimen DesignThe underlying hypothesis of this work is that the material modulus of honeycomb structures with a certain cell geometry will converge to a single value as the number of unit cells approaches “infinity”. Further, this limiting value of the material modulus will be the same for all honeycomb structures regardless of the cell geometry.

Further, this limiting value of the material modulus will be the same for all honeycomb structures regardless of the cell geometry. To test these hypotheses, twenty-five specimens, varying in the number of unit cells, were designed for each geometry group: regular hexagon, square, and equilateral triangle, which are the only three regular two-dimensional space-filling shapes.Specimens with the regular hexagon geometry had number of cells ranging from 5 to 105, specimens with the square geometry from 16 to 256 cells, and specimens with the equilateral triangle geometry had the number of unit cells ranging from 28 to 460 cells. All of the sample structures were designed using SolidWorksTM with a wall thickness (t) of 0.032", unit length (l) of 0.18", and base width (b) of 1". Three representative CAD models with dimensions for each shape are shown in Figures 3a-c.(a)(b)(c)Figure 3.

Three representative CAD models with dimensions for each shape are shown in Figures 3a-c.(a)(b)(c)Figure 3. Representative CAD models showing key parameters used in the design for (a) hexagonal, (b) square and (c) triangular honeycomb structures. Note: designs not to scale in above image with respect to each otherManufacturing and InspectionThe specimens were manufactured using the FDM process on a Stratasys © FortusTM 400mc machine with ABS-M30TM material. The samples were manufactured in three separate but successive builds on the same machine, using material from a single ABS canister, with the same tip and no re-calibrations performed between builds, to minimize process-induced variation.2151The printing parameters used were a layer thickness of 0.010" and a contour width of 0.016", which resulted in 100 layers and a wall thickness of two contours for each specimen.

The screenshot of how the filaments were laid out by Stratasys’s InsightTM software, and the corresponding printed specimens are shown in Figure 4. A total of 75 specimens were manufactured: 25 per shape. Each of these 75 specimens was unique – in other words, no replications were conducted since the intent of this work was to establish trends in the three shapes as a function of the size of the specimen. These specimens are shown in Figure 5.(a)(b)(c) Figure 4. Toolpath (left) and as-printed (right) layers for (a) hexagonal, (b) square and (c) triangular honeycomb structures2152A typical industrial FDM process is often capable to dimensional tolerances of approximately 0.005". As a result, for honeycomb structures with small thickness values (0.010-0.070"), this can result in a significantly large percentage of dimensional error.

As a result, for honeycomb structures with small thickness values (0.010-0.070"), this can result in a significantly large percentage of dimensional error. In order to verify the dimensions of these printed structures, four additional specimens from each shape were chosen as representative specimens to be scanned using blue light inspection technique with COMET LƎDTM. Thickness and length measurements were made with Geomagic ControlTM. The dimensions were measured at multiple locations including both edge and center locations at the top and bottom of the specimens. The findings of this study are summarized in Table 1; details of measurements made and the associated statistical analyses are reported in Appendix A towards the end of this paper. As the results in Table 1 show, the average thickness variation was found to be 0.005" in line with expectations. These values were used in the analytical calculations discussed later in the paper.Figure 5.

These values were used in the analytical calculations discussed later in the paper.Figure 5. The 75 specimens used in this study – 25 for each of the three shapes: honeycomb (left), square (center), and triangle (right)Table 1. Summary of the analysis of the thickness and length of the printed specimensHexagon Square TriangleThickness (designed 0.032") Standard deviation (in) 0.00126 0.00152 0.00155Mean (in)0.0315 0.0322 0.0316Length (designed 0.18") Standard deviation (in) 0.00018 0.00030 0.00020Mean (in)0.1800 0.1802 0.1802Mechanical TestingAll 75 specimens were tested on an Instron 8801 mechanical testing system operated in compression mode under displacement-control, with a displacement rate of 0.2" per minute, following experimental methods described in previous work [12].

All testing was completed without any changes to the setup of the machine or repositioning of the compression platens.2153Representative test results from hexagonal, square and triangular honeycomb specimens are shown in Figure 6. The red markers on the graphs correspond to the images captured from a concurrently running video and show how localized deformation in the specimen (encircled in the photographs) relates to observed global deformation. The observed failure mechanism appears consistent with the “plastic hinge” mechanism [13], which is expected for a plastic material such as ABS [21]. Further discussion of the failure mechanism is beyond the scope of the current work, which is concerned with estimation of the modulus only.(a)(b)2154(c)Figure 6. Load displacement graph under compression for (a) hexagonal, (b) square and (c) triangular honeycomb specimens.

Load displacement graph under compression for (a) hexagonal, (b) square and (c) triangular honeycomb specimens. The photographs correspond to the red markers on the graph.AnalysisThe main goal of this approach is to establish a material modulus that is both shape and size independent. To enable this, measured load-displacement responses from these different honeycomb structures need to first be translated into a stress-strain plot and then the modulus calculated. For this purpose, the member modeling method discussed previously and elaborated in previous work [12] was adapted for use here, with two key differences discussed below.A recent paper by Malek and Gibson [23] corrected the commonly used equations for honeycomb modulus estimation [13] by redefining the length of the connecting members. This corrected length was used in all calculations in this work.

This corrected length was used in all calculations in this work. For all but the thinnest members, this correction can have a significant effect on modulus calculations (a 10-20% increase). Secondly, this work sought to relate load and displacement at the specimen level to the values at the level of an individual beam. For hexagonal honeycombs loaded in compression, the beam members that are neither parallel nor normal to the applied load (“diagonal” beam members) are the most strained, as shown in Figure 7. This was done to minimize errors associated with defining a specimen area and length for normalization in converting load-displacement to stress-strain. Thus, the total load at the specimen level was divided by the number of beam members that run across the specimen to arrive at a load per beam member. Similarly, the total displacement was divided by the number of diagonal beams to arrive at a displacement per beam member.2155Figure 7.

Similarly, the total displacement was divided by the number of diagonal beams to arrive at a displacement per beam member.2155Figure 7. Simulation result of a hexagonal honeycomb specimen under compression showing diagonal beam elements taking up the bulk of the elastic strainUsing this method eliminates the need for defining an effective modulus and only invokes standard beam theory to solve for a material modulus Es [12, 13]. Thus for a beam member with thickness t, effective length lb, width b, the material modulus can be described in terms of experimentally measured, beam-level load P and displacement δ, per the equations below:For hexagonal honeycombs:32Plb(sin 30°) 3Es =δbtFor square honeycombs:(2)(3)PlbEs =δbtFor triangular honeycombs:Plb(cos 30°)2(4)Es =δbtThe effective length lb follows from Malek and Gibson for hexagonal honeycombs [23] and similar equations can be derived for square and triangular honeycombs.

Different approaches were evaluated to quantify the effect of the number of cells on elastic modulus, including using the number of cells in each direction as shown in Figure 1, where the effect of cell quantity on modulus is not the same in both directions. As a result, a new parameter that accounted for both variables together was introduced, which segregates cells in a specimen into partially and fully constrained cells, as shown in Figure 8. This then enables a calculation of a percentage of (fully) constrained cells. The motivation for this approach derives from the fact that analytical equations assume no boundary effects and an infinite number of cells in all directions [13]. Thus, a specimen with 100% percentage of its cells fully constrained would represent the theoretical ideal.2156Figure 8.

Thus, a specimen with 100% percentage of its cells fully constrained would represent the theoretical ideal.2156Figure 8. Definition of partially and fully constrained cells and the directionsUsing the aforementioned equations, measured load and displacement responses from the compression tests and the dimensions measured using optical scanning, elastic modulus values were computed for each of the 75 samples tested. Each of these data points is plotted in Figure 10 for hexagonal, square and triangular cells. Whereas the general trend is that the estimated material modulus increases as the percentage of constrained cells increases, other interesting trends are also observed and indicated by the alphabetic grouping. These trends, it is hypothesized, emerge as a result of differing edge behaviors in the experiments conducted.

These trends, it is hypothesized, emerge as a result of differing edge behaviors in the experiments conducted. As shown in the schematic in Figure 9, cells on the top and bottom are constrained by friction with the compression platens, the cells on the left and right are partially constrained.Figure 9. Frictional constraint and free surface (partially constrained cells) create boundary conditions that are not part of the theoretical model and each has a different effect.2157(a)(b)(c) Figure 10. Load displacement graph and snapshots in the deformation process under compression for (a) hexagonal, (b) square and (c) triangular honeycomb specimens.2158In Figure 10 (a), the specimens in group A had 2 transverse cells, and the specimens in group B had 2 axial cells. Thus the majority of cells in these groups were in contact with the platens (frictional constraint) or had free surfaces (partially constrained).

Thus the majority of cells in these groups were in contact with the platens (frictional constraint) or had free surfaces (partially constrained). For specimens in group C, the correlation between increasing percentage of constrained cells and modulus is fairly robust. Similar to the hexagonal specimens, the results in Figure 10 (b) showed a divergence in the values of the material modulus for the square specimens that had low percentage of constrained cells. Group A contained all the specimens that had only 4 axial cells, while group B contained all the specimens that had 4 transverse cells. Group C contained all of the specimens that had 7 axial cells, and group D contained all the specimens that had 7 transverse cells. The identical trend was observed for the triangular honeycombs as well in Figure 10 (c). Group A had 4 axial cells, B had 4 transverse cells, C had 8 axial cells and D had 6 transverse cells.

The identical trend was observed for the triangular honeycombs as well in Figure 10 (c). Group A had 4 axial cells, B had 4 transverse cells, C had 8 axial cells and D had 6 transverse cells. The raw data that led to the estimation of these modulus values is compiled in Appendix B.ResultsThe trends in Figure 10 are indicative of the core of the challenge this work seeks to address: any measurements of modulus made by characterizing honeycomb structures are strongly dependent on both shape (hexagon, square or triangle) and size (percentage of constrained cells). The hypothesis of this work is that with a member modeling approach that is inherently shape independent, a saturated value of modulus may be extracted for large enough specimens. To determine the validity of this hypothesis with the data in Figure 10, the 95% confidence interval approach was used.

To determine the validity of this hypothesis with the data in Figure 10, the 95% confidence interval approach was used. The data was compartmentalized in intervals of reducing size (with a decrement of one for each sample, eliminating the point corresponding to the lowest percentage of constrained cells for each iteration) until all the material modulus values extracted from those specimens fell within the range of the 95% confidence interval around their mean. On arriving at this value, it was concluded that the mean of those values could be taken as the representative material modulus. Applying this method to all three shapes resulted in the modulus values in Table 2, which are listed along with the percentage constrained cell values for the specimens that were used to compute the mean modulus.Table 2.

Summary of the analysis of the thickness and length of the printed specimensRegular hexagon Square Equilateral triangle% Constrained Cell Values for Specimen Range in 95% Interval 55-64 72-77 68-74Extracted material modulus (ksi)273 248 262Difference from Hexagonal Honeycomb - 9.2% 4.2%The agreement of the material modulus extracted from characterization of square and triangular specimens with hexagonal ones is under 10% is a key accomplishment of this work. This difference can be explained by the limited sample size and by the fact that theory assumes cell shapes with sharp corners whereas in reality these have some curvature to them. For a sanity check, bulk ABS values are often reported in the 260-360 ksi range [21], and Stratasys© reports values of 310-320 ksi for specimens made with the FDM process [23] – both these numbers are2159close to the values reported in Table 2.

In addition to the causes for difference described above, the beam geometries are also not perfectly rectangular as theory assumes, which could be an additional reason why values are different from bulk values.ConclusionsThe following conclusions may be drawn from this work:1. This work confirms experimentally previously documented observations that modulus estimations for cellular structures are highly dependent on the size of the sample used for characterization.2. This work also shows how using a beam-modeling approach, with the correct definition of beam length and calculating beam-level load and displacement from experimental measurements, can be used to extract a material modulus that is shape independent.3.

Using a metric of percentage of constrained cells and measuring modulus enables the monitoring of these trends and further find near-saturation values that approach size independence.Extracted values of modulus using this approach from different shapes and sizes are within 10% of each other – and may be used with greater confidence for non-uniform geometries such as Voronoi cells, as long as they have the same inherent beam-level structure that was characterized to extract the modulus.

This approach can be extended to cellular structures of different materials and AM processes, and it is likely to be successful for failure prediction as well.AcknowledgementThis effort was performed through the National Center for Defense Manufacturing and Machining under the America Makes Program entitled “A Non-Empirical Predictive Model for Additively Manufactured Lattice Structures” and is based on research sponsored by the Air Force Research Laboratory under agreement number FA8650-12-2-7230. The U.S. Government is authorized to reproduce and distribute reprints for Governmental purposes notwithstanding any copyright notation thereon.DisclaimerThe views and conclusions contained in this document are those of the authors and should not be interpreted as necessarily representing the official policies, either expressed or implied, of the Government. Distribution authorized to U.S. Government Agencies and America Makes Members; Critical Technology.

Distribution authorized to U.S. Government Agencies and America Makes Members; Critical Technology. Other requests for this document shall be referred to AFRL/RXMS, Wright-Patterson Air Force Base, OH 45433-7750.2160Appendix A: Dimensional DataThis appendix provides data on the measurements made using blue light scanning to extract true, as-printed dimensions. These dimensions (and not as-designed values) were used in the analysis.Figure A1. Designed and scanned data superimposed, measurements made with Geomagic ControlTM2161Figure A2. Analysis of the thickness of the regular hexagon specimens. Of note is the difference in the means comparisons for bottom/top and center/edge.Figure A3. Analysis of the length of the regular hexagon specimens2162Figure A4. Analysis of the thickness of the square specimens. Of note is the difference in the means comparisons for bottom/top and center/edge.Figure A5. Analysis of the length of the square specimens2163Figure A6.

Of note is the difference in the means comparisons for bottom/top and center/edge.Figure A5. Analysis of the length of the square specimens2163Figure A6. Analysis of the thickness of the equilateral triangle specimens. Of note is the difference in the means comparisons for bottom/top and center/edge.Figure A7. Analysis of the length of the equilateral triangle specimens.2164Appendix B: Compression Test DataThis appendix provides the load-displacement data for all 75 samples tested – to make viewing easier, each shape is broken down into 3 categories corresponding to the percentage of constrained cells in those specimen designs.Figure B1. Graphs of load vs. position of regular hexagon specimens that had 0-20% of constrained cells.Figure B2. Graphs of load vs. position of regular hexagon specimens that had 21-40% of constrained cells.Figure B3. Graphs of load vs. position of regular hexagon specimens that had more than 40% of constrained cells.2165Figure B4.

Graphs of load vs. position of regular hexagon specimens that had more than 40% of constrained cells.2165Figure B4. Graphs of load vs. position of square specimens that had 20-40% of constrained cells.Figure B5. Graphs of load vs. position of square specimens that had 41-60% of constrained cells.Figure B6. Graphs of load vs. position of square specimens that had more than 60% of constrained cells.2166Figure B7. Graphs of load vs. position of equilateral triangle specimens that had 20-40% of constrained cells.Figure B8. Graphs of load vs. position of equilateral triangle specimens that had 41-60% of constrained cells.Figure B9. Graphs of load vs. position of equilateral triangle specimens that had more than 60% of constrained cells.2167References1. C. Neff, N. Hopkinson, N.B. Crane, "Selective Laser Sintering of Diamond Lattice Structures: Experimental Results and FEA Model Comparison," 2015 Solid Freeform Fabrication Symposium2. M. Jamshidinia, L. Wang, W.

Crane, "Selective Laser Sintering of Diamond Lattice Structures: Experimental Results and FEA Model Comparison," 2015 Solid Freeform Fabrication Symposium2. M. Jamshidinia, L. Wang, W. Tong, and R. Kovacevic. "The bio-compatible dental implant designed by using non-stochastic porosity produced by Electron Beam Melting®(EBM)," Journal of Materials Processing Technology214, no. 8 (2014): 1728-17393. S. Park, D.W. Rosen, C.E. Duty, "Comparing Mechanical and Geometrical Properties of Lattice Structure Fabricated using Electron Beam Melting", 2014 Solid Freeform Fabrication Symposium4. D.M. Correa, T. Klatt, S. Cortes, M. Haberman, D. Kovar, C. Seepersad, "Negative stiffness honeycombs for recoverable shock isolation," Rapid Prototyping Journal, 2015, 21(2), pp.193-2005. C. Yan, L. Hao, A. Hussein, P. Young, and D. Raymont. "Advanced lightweight 316L stainless steel cellular lattice structures fabricated via selective laser melting," Materials & Design 55 (2014): 533-5416. S.

Hao, A. Hussein, P. Young, and D. Raymont. "Advanced lightweight 316L stainless steel cellular lattice structures fabricated via selective laser melting," Materials & Design 55 (2014): 533-5416. S. Didam, B. Eidel, A. Ohrndorf, H.‐J. Christ. "M echanical Analysis of M etallic SLM ‐ Lattices on Small Scales: Finite Element Simulations versus Experiments," PAMM 15.1 (2015): 189-1907. P. Zhang, J. Toman, Y. Yu, E. Biyikli, M. Kirca, M. Chmielus, and A.C. To. "Efficient design-optimization of variable-density hexagonal cellular structure by additive manufacturing: theory and validation," Journal of Manufacturing Science and Engineering 137, no. 2, 20158. M. Mazur, M. Leary, S. Sun, M. Vcelka, D. Shidid, M. Brandt. "Deformation and failure behaviour of Ti-6Al-4V lattice structures manufactured by selective laser melting (SLM)," The International Journal of Advanced Manufacturing Technology 84.5, 2016, 1391-1411 9. R. Gümrük, R.A.W.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDETERMINATION OF INPUT LASER ENERGY FOR MELTING POWDER LAYERS OF VARIOUS THICKNESSES IN HIGH-SPEED PBF-LB/P USING NEAR- INFRARED LASER AND ABSORBENTYuki Yamauchi*, Takashi Kigure*†, and Toshiki Niino†Tokyo Metropolitan Industrial Technology Research Institute, Aomi, Koto-ku, Tokyo, 135- 0064 †Institute of Industrial Science, the University of Tokyo, Komaba, Meguro-ku, Tokyo, 153- 8505 AbstractThe rate of production of PBF-LB/P can be increased by increasing the layer thickness. However, this reduces the part resolution in the stacking direction. To obtain both a high rate of production and high part resolution, layer thickness adjustment in accordance with part geometry can be effective. Optimizing the input laser energy with respect to the layer thickness ensures sufficient melting and part strength.

Optimizing the input laser energy with respect to the layer thickness ensures sufficient melting and part strength. According to previous studies, the use of a near- infrared laser and absorbent can increase penetration depth or depth of fusion. However, the optical properties of the powder bed can vary significantly depending on the layer thickness, and, therefore, the input energy that actually contributes to melting also changes with layer thickness.

This study proposes a method for determining the input laser energy for various layer thickness without trial and error by estimating the amount of energy required to melt the powder layer while accounting for the optical properties of the bed.IntroductionPowder bed fusion with laser beam and polymeric material (PBF-LB/P), an additive manufacturing technology using polymeric materials with relatively high productivity and high part reliability, is increasingly being adopted as a production method for end-use parts; it reduces the number of assembly processes and reproduces complex shapes [1]. However, the productivity of PBF-LB/P is significantly lower than that of injection molding, the process which is currently used. This results in high component costs, hindering its widespread use. To improve the productivity of PBF-LB/P, it is necessary to increase the build process speed.

This results in high component costs, hindering its widespread use. To improve the productivity of PBF-LB/P, it is necessary to increase the build process speed. This can be achieved by shortening the time required to create each layer[2] or by reducing the number of layers per part. The number of layers required to build each part can be reduced by increasing the layer pitch (nominal layer thickness) [3][4][5][6], although the layer pitch is closely related to the part resolution in the stacking direction. As proposed here (see Figure 1), by adjusting the stacking pitch according to the part shape, the number of layers can be reduced while maintaining the resolution of the part. As the layer pitch changes, the amount of powder used in each layer and the thickness of the powder layer change. As the powder layer thickness increases, more energy is required to melt it and it becomes more difficult to connect it with the layer underneath [7][8].

As the powder layer thickness increases, more energy is required to melt it and it becomes more difficult to connect it with the layer underneath [7][8]. Since the melting condition of the powder and the adhesion between the layers directly affect the mechanical properties of the built part[9], an appropriate amount of laser energy should be supplied to sufficiently melt the powder layers of various thicknesses. Also, high energy input to melt thick (deep) powder layers may cause thermal degradation of the polymer near the surface of the powder layer[10][11][12][13]. This is especially important for light that is highly absorbed, such as a CO2 laser, which has low penetration through the powder layer[14][15]. Previous research by the authors has shown that increasing the laser transmittance increases the melting depth while minimizing thermal degradation[16].

Previous research by the authors has shown that increasing the laser transmittance increases the melting depth while minimizing thermal degradation[16]. Therefore,46for processes employing variable powder layer thickness, an increase in the transmittance of the laser is an effective solution. However, the method employed by the authors to adjust the transmittance of a near-infrared laser – the use of an absorber – results in a large loss of energy due to diffuse reflection, which varies depending on the powder layer thickness. Therefore, the layer thickness and optical properties must be considered to systematically and efficiently determine the appropriate input laser energy. The authors have modified EMR (Energy Melt Ratio which is the ratio of the amount of energy required to melt the powder material and the amount of input laser energy, proposed by Starr et al.

The authors have modified EMR (Energy Melt Ratio which is the ratio of the amount of energy required to melt the powder material and the amount of input laser energy, proposed by Starr et al. [9]) taking into account the loss of energy due to the reflection and transmission of laser light in the powder bed, and have attempted to quantitatively determine the amount of input laser energy [17]. In addition, the relationship between the modified EMR and the part density, which is strongly correlated with its mechanical properties, was investigated and was found to be good and stable regardless of the various conditions, such as the optical properties of the powder bed. In other words, it is possible to determine the input laser energy for various layer pitches by maintaining the modified EMR value proposed by the authors.

In other words, it is possible to determine the input laser energy for various layer pitches by maintaining the modified EMR value proposed by the authors. In this study, we determined the amount of input laser energy for a given and fixed modified EMR value for various layer pitches without trial and error, and investigated the relationship between the layer pitch and the mechanical properties of the built part.Figure 1 Schematic of layer pitch adjustment in a part47Methods and MaterialsMethodologyFor melting thick powder layers without thermal decomposition near the surface, the depth up to which the incident laser beam reaches, i.e., the penetration depth, should be larger than that of a CO2 laser. In this study, to increase the penetration depth, we used a near-infrared laser, which permeates most polymeric materials and added an absorber to the powder material for absorbing near-infrared light, as mentioned earlier.

By keeping the modified EMR value for each layer pitch constant as a reference, the quantitative input laser energy can be matched even if the layer pitch changes. The modified EMR values used in this study were determined based on the relationship between various modified EMR values and the part density which is closely related to melting. Tensile tests were conducted on the built specimens to check whether the part strength could be maintained constant while changing the layer pitch and unifying the modified EMR values.

Tensile tests were conducted on the built specimens to check whether the part strength could be maintained constant while changing the layer pitch and unifying the modified EMR values. In this study, specifically, the tensile strength in the stacking direction, which is strongly affected by an increase in layer pitch, was evaluated.𝐸𝐸𝐸𝐸𝐸𝐸𝑚𝑚 = , , ,,, and,,,, laser power, scan spacing,𝐷𝐷𝑝𝑝ℎ𝑓𝑓𝐸𝐸𝑧𝑧𝐸𝐸0𝑦𝑦𝑣𝑣𝑃𝑃𝑧𝑧𝑧𝑧MaterialsPA11 (ASPEX-FPA, Aspect Inc.) was used as the powder material, and 0.4 wt% of acoloring agent (Nubian Black PC-0870, Orient Chemical Industry Co., Ltd.) was added in powder form to the PA11 powder to assist laser light absorption in the near-infrared region. for this material, were referenced from the specifications of a similar,,,, andmaterial [18] and were set to 2.09 J/g K, 200°C, 0.45 g/cm3, and 83.7 J/g, respectively.

for this material, were referenced from the specifications of a similar,,,, andmaterial [18] and were set to 2.09 J/g K, 200°C, 0.45 g/cm3, and 83.7 J/g, respectively. 𝐶𝐶𝑝𝑝 Optical properties measurement𝑇𝑇𝑚𝑚𝜌𝜌𝑝𝑝 𝑇𝑇𝑏𝑏ℎ𝑓𝑓The optical properties of the powder material used in this study, including,, and, were measured for the modified EMR calculation in same manner as proposed in the 𝐸𝐸𝑧𝑧 authors’ previous report [8] [17]. The thickness of the powder layer was calculated from the 𝐷𝐷𝑝𝑝 weight of the powder filled in the container and the bulk density of the powder. The container consisted of quartz glass and a stainless steel plate with a 50 mm diameter hole. The volume of the container and the amount of powder filled were adjusted by changing the thickness of the stainless steel plate. The powder layer formed on the container was irradiated by a laser beam with a diameter of 1.5 mm and a wavelength of 1.06 µm.

The powder layer formed on the container was irradiated by a laser beam with a diameter of 1.5 mm and a wavelength of 1.06 µm. The light passed through the powder layer and the bottom of the container (quartz glass), and the transmitted light was measured with a power meter. Transmittance is the ratio of the energy transmitted through the powder layer and the incident energy. Diffuse reflectance was measured simultaneously with transmittance measurements using a barium sulfate-coated integrating sphere and the same light source, container, and powder material as for transmittance measurements. To𝐸𝐸048compensate for the loss of energy caused by the integrating sphere, the diffuse reflectance of the powder layer was determined from the ratio to a reference specimen of known reflectance (99.6%).Powder layer thicknessIn this study, we make a clear distinction between layer pitch and powder layerthickness.

Because the low bulk density of the powder results in the formation of a powder layer thicker than the layer pitch in the next layer due to consolidation by melting and flow [19][20]. Consolidation is considered in the powder layer thickness equation below:as shown in the𝑧𝑧𝑧𝑧 ≅ 𝑧𝑧𝑖𝑖 𝜌𝜌𝑚𝑚 where, 𝜌𝜌𝑝𝑝 layer thickness can be greater than this equation for the first few layers, as shown below. In the melting process of the first layer, the laser beam is irradiated into a very thick powder layer, since there is no previously processed layer. Therefore, the melting depth and consolidation depth can be greater, leading to recoating of the thicker powder in the next layer. Therefore, this equation is valid when the recoating process has been done with several layers.andare layer pitch and density of the melting region, respectively.

Therefore, this equation is valid when the recoating process has been done with several layers.andare layer pitch and density of the melting region, respectively. Note that the (2)𝑧𝑧𝑖𝑖𝜌𝜌𝑚𝑚Specimen preparationA laser sintering machine (RaFaEl300F, Aspect Inc.) equipped with a near-infrared laser (wavelength = 1.06 µm) was used to build the specimens. Specimens for the density measurements were 10 × 10 × 5 mm in size, with a layer pitch of 100 µm and a modified EMR ranging from 1.2 to 3.2. The process parameters are shown in Table 1. The powder bed temperature was set at 140°C, which is lower than the process window. According to the authors' previous study, when the powder bed temperature is low, especially below the process window, bonding between the processing layer and the previously processed layer becomes difficult. This means that we can easily observe the effect of changes in layer pitch on the bonding process under the bed temperature below the process window.

This means that we can easily observe the effect of changes in layer pitch on the bonding process under the bed temperature below the process window. Therefore, the powder bed temperature was set below the process window in this study.For a high-speed processing, the laser beam is scanned as fast as possible, and then the high laser power is required to fill the input energy for appropriate melting.

However, the maximum laser power of the machine used in this study is not sufficient, so a slow scanning speed was used.Contour scanning was not performed for the following reasons: it does not have a large effect on the part density of the fabricated part, and for the tensile test specimen, the surface was removed by milling or polishing before the tensile test.49Table 1 Process parametersBased on the results of the density evaluation described below, specimens for tensiletesting were fabricated for a layer pitch range of 100-300 μm with a modified EMR of approximately 1.7, the minimum value at which the part density saturates.Part densityA relative density of the built part was obtained to know the melt state. The relativedensity is the ratio of the density of the part to the true density of the material used; the absorbent is not included in the calculation because its concentration is negligibly small.

The relativedensity is the ratio of the density of the part to the true density of the material used; the absorbent is not included in the calculation because its concentration is negligibly small. The density of the built part was measured based on the Archimedes method.Tensile testTensile tests were performed using a universal testing machine (INSTRON 3366, 10kN load cell). Tensile specimens were of standardized geometry and 1/2 scale, that is 75 × 10 × 2 mm. The longitudinal side of the specimen was oriented in the stacking direction (z axis). The crosshead travel speed was set at 2 mm/min according to the specimen size. To suppress the influence of surface roughness on the tensile test result, the specimen surface was removed by milling or polishing before the test as mentioned above.Results and DiscussionOptical propertiesFigure 2 shows the relationship between the powder layer thickness and transmittance.

Assuming that the attenuation of the laser light incident on the powder layer is based on the Beer-Lambert law [21], the measured values were approximated by an exponential curve. The penetration depth obtained from the approximate curve was 225 µm. According to previous reports, the penetration depth of CO2 lasers is 80 to 100 µm [14][15], which is less than half as deep as that in this study.Figure 3 shows the relationship between the powder layer thickness and diffuse reflectance. The reflectance increases with the powder layer thickness and tends to saturate above 300µm. Assuming that the diffuse reflection in this study follows the Kubelka-Munk theory [22], an approximate curve can be obtained as shown in the figure from which the diffuse reflectance at a given thickness can be obtained.

For all thicknesses, it can be seen that the diffuse reflectance is very high for the method used in this study compared to that of the CO2 laser, which has been reported to be less than 5% [14][23], indicating that the energy loss due to reflection must be considered. Additionally, specular reflection was also obtained by the previously reported method [17] and was approximately 8%.50100% e c n a t t im s n a r T10100.2 0.6 0.4 Thickness of powder layer mm0.8Figure 2 Relationship between powder layer thickness and transmittance. Error bars were obtained from the maximum and minimum values of the experiment for n = 5.706050% e c n a t c e l f e R40302010000.4 Thickness of powder layer mm0.20.60.8Figure 3 Relationship between powder layer thickness and diffuse reflectance.

Error bars were obtained from the maximum and minimum values of the experiment for n = 5.Part densityFigure 4 shows the relationship between the modified EMR and the relative density of the built specimens. The results show that the relative density of the parts saturates when the modified EMR is greater than 1.5. Theoretically, the powder should melt at a modified EMR of 1, but because of the energy required for coalescence and flow, a modified EMR greater than 1 is required to increase the density.5110095% y t i s n e d e v i t a l e R9085807570656001.02.0 Modified EMR3.04.0Figure 4 Relationship between modified EMR and relative density of built specimens. Error bars were obtained from the maximum and minimum values of the experiment at n = 5.Tensile testFigure 5 shows the relationship between the layer pitch and ultimate tensile strength (UTS).

Error bars were obtained from the maximum and minimum values of the experiment at n = 5.Tensile testFigure 5 shows the relationship between the layer pitch and ultimate tensile strength (UTS). For layer pitch values from 100 to 250 µm, the UTS remains almost flat at about 35 MPa, showing that the same modified EMR value can be used to build parts with similar strengths. When the layer pitch is 300 µm, the UTS is about 30 MPa, which is 15% lower compared to the other values. With respect to the Young's modulus result, as shown in Figure 6, it appears to be generally close to the UTS result. The decrease in Young's modulus begins above 250 µm pitch, which is slightly different from the UTS result. There is also almost the same trend in elongation at break as the UTS result, except for the 150 µm pitch (see Figure 7). Figure 8 shows the cross sections of the samples with different layer pitches. Voids can be seen for all pitches.

Figure 8 shows the cross sections of the samples with different layer pitches. Voids can be seen for all pitches. However, the 300 µm pitch specimen has more regularly positioned voids which might have reduced its strength in the stacking direction.The result of the tensile test and void formation are discussed below. As stated earlier, the laser penetration depth of the powder bed in this study was 225 µm. But, for each layer, the powder layer thickness in the processed area is approximately 2.3 times the layer pitch, based on the relationship between the layer pitch and the density of the specimen as given by equation (2). Therefore, when the layer pitch is 100 µm or greater, the powder layer thickness is greater than the laser penetration depth. Powder in the regions deeper than the penetration depth is heated and melted mainly by heat transfer from the regions where the laser beam penetrates and heats directly.

Powder in the regions deeper than the penetration depth is heated and melted mainly by heat transfer from the regions where the laser beam penetrates and heats directly. As the layer pitch is increased, the powder in the deeper, impenetrable regions is not sufficiently melted because not enough energy is transferred. This leads to a lack of bonding between the layers. On the other hand, if the input laser energy is increased to melt the deep- seated regions sufficiently, the surface of the powder bed could exceed the thermal decomposition temperature. Therefore, depending on the depth of the powder layer and light penetration, it can be assumed that thermal decomposition at the surface of the powder bed and insufficient melting at the bottom of the powder layer, both, will occur in the same layer as shown in Figure 9.

This means that even though it may be necessary to increase the input energy to eliminate insufficient melting, it will not be possible because it would lead to excessive thermal decomposition on the surface. Voids of different shapes are present in the 300 µm pitch52in Figure 8, implying that they were generated by different phenomena. If the previous assumption is correct, some voids are caused by insufficient melting and the others by thermal decomposition.As can be understood from the previous discussion, it is necessary to make the temperature difference between the surface and bottom of the powder layer as small as possible in order to further increase the layer pitch. To achieve this, it should either be possible to use a low energy input laser or to increase the penetration depth of the powder bed.

To achieve this, it should either be possible to use a low energy input laser or to increase the penetration depth of the powder bed. The first option can be realized by increasing the powder bed temperature, but there is an upper limit to this, and the degradation of the material due to temperature must be fully considered. The second option can be implemented by using a small amount of absorbent to increase the penetration depth. Previous studies by the authors show that the penetration depth of the powder bed can be controlled by a calibration curve, which is a relationship between optical properties (absorbance) and absorbent content (concentration) based on the Beer-Lambert law. For example, in this study, when the absorbent content is reduced to 1/2, the penetration depth is expected to increase to 300 µm.

For example, in this study, when the absorbent content is reduced to 1/2, the penetration depth is expected to increase to 300 µm. But this also means lower absorption, and so the actual input energy required for processing will be greater.In addition, the UTS result in this study appears to be slightly lower than other commercially available parts using the same material. To obtain higher UTS, higher interlayer adhesion is required, which might be obtained under higher M-EMR. Therefore, optimizing the M-EMR in this way is important and should be carried out in future studies.Finally, we tested only one type of material, absorbent and absorbent content in this study. We also need to validate in future study whether the method proposed in this study can be applied to other material.4035a P M S T U302520150100200 Layer pitch µm300400Figure 5 Relationship between layer pitch and UTS under modified EMR of approximately 1.7.

Error bars were obtained from the maximum and minimum values of the experiment for n = 5.5317001600a P M1500s u l u d o m1400s g n u o Y'13001200110010000100200 Layer pitch µm300400Figure 6 Relationship between layer pitch and Young’s modulus under modified EMR of approximately 1.7. Error bars were obtained from the maximum and minimum values of the experiment for n = 5.1412% k a e r bt a108n o i t a g n o l E64200100200 Layer pitch µm300400Figure 7 Relationship between layer pitch and Elongation at break under modified EMR of approximately 1.7. Error bars were obtained from the maximum and minimum values of the experiment for n = 5.54Figure 8 Cross sections of build specimens observed under a microscope with transmitted light. One of the voids in each image is indicated by an arrow.

One of the voids in each image is indicated by an arrow. 20 µm thick specimens were cut for observation from cross sections of tensile specimens using a microtome.Figure 9 Insufficient melting and thermal decomposition in the same layer. Voiding generated by thermal decomposition results in a spherical shape, while those caused by insufficient melting are distorted and sharply pointed.55ConclusionIn this study, in order to efficiently and systematically determine the input energy of the laser for various layer pitches, the energy absorbed by the powder was quantitatively determined from its optical properties, and the relationship between the modified EMR, which is the ratio of the amount of energy required to melt the powder material and the amount of input laser energy, and the part density was obtained. For the material used in this study, it was found that high-density parts could be built with a modified EMR of 1.5 or higher.

For the material used in this study, it was found that high-density parts could be built with a modified EMR of 1.5 or higher. Furthermore, the mechanical properties of specimens built with a fixed modified EMR and different layer pitches were evaluated. The results indicate that similar mechanical properties can be obtained when the stacking pitch is 250 µm or less. The results of this study suggest that if the input laser energy is controlled so that the modified EMR becomes a set value for a certain layer pitch or less, it is possible to set the desired layer pitch while retaining the mechanical properties, thereby enabling efficient fabrication and maintaining part resolution.AcknowledgementThis research was supported by METI R&D Support Program for Growth-orientedTechnology SMEs Grant, Number JPJ005698.Reference[1][2][3]T.D. Ngo, A. Kashani, G. Imbalzano, K.T.Q. Nguyen, D.

termination of the Optimum Joint Design for LENS Fabricated Ti6Al4V and Ti6Al4V/TiC Dual-Material StructuresObielodan J.O1 and Stucker B.E2*1Mechanical and Aerospace Engineering Department, Utah State University, Logan, UT 843222Industrial Engineering Department, University of Louisville, Louisville, KY 40292AbstractJoints between dissimilar material systems made using laser metal deposition processeshave been investigated. The fusion of materials with different physical properties and chemicalcompositions under high laser power often results in defects at the joints. Although somesolutions have been suggested in previous work for defect-free fabrications, most of the jointsstudied have been characterized using qualitative techniques only. Quantitative study isimperative for predicting the mechanical behavior of fabricated structures for real lifeapplications.

Quantitative study isimperative for predicting the mechanical behavior of fabricated structures for real lifeapplications. In this work, tensile and flexural specimens made of different Ti6Al4V andTi6Al4V/10%TiC dual-material transition joint designs were fabricated using laser engineerednet shaping (LENS) and tested. It was found that transition joint design has a significant effecton the tensile strengths of dual-material structures.1.IntroductionThe capabilities of additive manufacturing (AM) technologies to fabricate complexgeometries have been widely demonstrated. AM has also been shown to enable the fabrication ofheterogeneous materials and microstructural mixes in a single component. Some technologieslike 3D printing [1-2], stereolithography [3-4], selective laser sintering [5], ultrasonicconsolidation [6-8], direct laser metal deposition processes [8-11] and others have been used todemonstrate the multi-material capabilities of AM.

The materials used range from polymers tometals and ceramics; and in some cases are process-specific. The chemical and physicalcompatibilities of the material systems are important factors that determine the qualities of thejoint between multiple materials. The use of difficult-to-join material systems is a challenge,however several solutions have been shown to enable the successful. Two examples are the useof gradient transitions from one material to the other [12] and the use of compatible intermediatematerials [7].Inter-material joint problems are common in fusion-based processes such as direct lasermetal deposition processes. The processes include laser engineered net shaping (LENS) and itsvariants, like direct light fabrication (DLF), epitaxial laser metal forming (E-LMF), laser directforming (LDF), laser rapid forming (LRF) and others.

Good selection of process parameters arerequired in order to achieve defect free component fabrication in all cases.LENS possesses the capabilities to fabricate fully dense structures using powdermaterials. It fabricates solid objects in a layer-wise fashion from computer aided design (CAD)models that are first numerically sliced to predetermined thickness. Each layer is fabricated bymelting powder materials delivered to the focus of the laser beam on a substrate that is mountedon an x-y stage. The stage moves in a raster fashion according to the tool paths generated usingthe sliced CAD models. The fabrication takes place under a controlled, inert atmosphere in aglove box. The process is schematically illustrated in Fig. 1. Some of the important processparameters are laser power, powder flow rate, layer thickness, hatch width, deposition speed andoxygen level in the glove box. The capabilities of the process for multi-material fabrication havebeen demonstrated.

The capabilities of the process for multi-material fabrication havebeen demonstrated. It is used for composite material fabrication [13-14], functionally gradientstructures [11-12], multi-materials processing [10] surface cladding for corrosion resistance [11],and biomedical applications [15-16].Figure 1: Schematic illustration of LENS deposition processMost of the earlier work on multi-material fabrication using direct laser metal depositionprocesses does not go beyond establishing successful fabrication of different material systemsthrough qualitative characterizations using microstructure studies. However, its potential forfabricating multi-material structures for load carrying applications will not be fully achievedwithout establishing the mechanical properties of transition joints between the materials.Dual-material minimum weight structure design [17-19] is one of the evolvingapplication areas of AM fabricated multi-material structures.

They are both geometrically andmaterially complex, and thus are difficult or impossible to fabricate using conventionalprocesses. They are made of multiple, thin members that are preferably made from light weightmaterials with high specific strength and stiffness. Such structures are readily applicable toaerospace and automotive industries, where there is continuous emphasis on higher strength andlower weight structures for improved fuel efficiency and performance. Figure 2 shows anexample of a complex minimum weight structure with members that could be made of differentmaterials based upon Michell theory [21]. In the illustrated structure, if the structure is pinned atpoints A and B and a load is placed at C, parallel to a line between A&B, as shown with thearrow, the outer member labeled D will be in pure compression, as well as all the inner membersthat join D tangentially.

Those inner network members that are perpendicular to D, and themember between A and C will be in pure tension. In order to optimize a structure to its fullestextent, the members in tension can be made of materials different from those in compression. Inthis case, the intersection between the tensile and compressive members and the design andstrength of these joints is of critical importance for the structure’s reliability and performance.ABDCFigure 2: A minimum weight structure design [21]Simplified minimum weight structures that are representative of the more complex designshown in Fig. 2 were designed based on maximum strength and maximum stiffness criteria [17]and fabricated. Figure 3 shows a free body diagram of a simplified structure design with oa, oband oc as compression members and ab and ac as tension members when subjected tocompressive load F with simple supports at b and c.

Given such a design withspan = L,applied force =,F2fF,where, f is the stress acting on the tension members at any point during loading, Table 1 showsthe load relationships existing in the structure members. Structures designed based on maximumstiffness criterion must satisfy the following strain ratio [17].2121CTTCCCTTCTEEEETable 1: Load and size relationship for minimum weight structureElementLengthForceCross-sectional areaoc2/L2/2/fF)2/(CFoa2/L2fF)/(CFac2/LfF2/)2/(TFFigure 3: Free body diagram of the dual-material minimum weight structure(i)This work is aimed at characterizing the failure characteristics and strengths of dual-material systems using different material transition designs at the joints between tensile andcompressive members.

Ti6Al4V/TiC composite and Ti6Al4V materials were used for the study.Different material transition joints were designed and tested for flexural and tensile strengths.Optimal designs were applied for the fabrication of dual-material minimum weight structures andtested.2Experimental ProceduresA laser engineered net shaping (LENS 750) machine made by Optomec Inc.,Albuquerque, USA, was used for this experiment. The machine uses a continuous waveND:YAG laser with a capacity of up to 400 Watts. The laser power used ranged from 200 to 270Watts. The machine has a dual powder feeder system that allows the simultaneous delivery oftwo different material mixtures. The powder is delivered by argon carrier gas to the focus of thelaser beam, and deposits were made on a 6mm thick commercially pure (CP) titanium substrate.The machine has a 3- axis motion system consisting of an x-y motion stage and a z-axis forintegrated laser and powder delivery system.

The oxygen level was maintained under 10 partsper million (ppm) in the glove box. Deposition layer were of 0.25mm thickness and 0.38mmhatch width for all fabrications.Spherical Ti6Al4V powder material of 125 – 210 microns diameter supplied byAdvanced Specialty Metals (ASM), New Hampshire, USA and TiC powder of 45 – 150 micronparticle size, supplied by Pacific Particulate Materials (PPM) Limited, Canada, were. Althoughthe recommended powder particle size for LENS fabrication is 45 – 150 microns diameter, theTi6Al4V size used was found to flow well through the delivery system. Two sets of dual-material specimens were fabricated for joint strength characterization. A set was fabricated forflexural strength determination and another set of corresponding joint designs for tensileproperties determination.In each test type, six different joint designs were fabricated with three replicates. Thedesigns, as shown in Fig.

Thedesigns, as shown in Fig. 4 below are: butt joint; gradient transition joint; interlocked materialjoint, randomly interlocked material joint, scarf joint and v-groove joint. The joints arerespectively labeled butt, gradient, interlock, random, scarf, and v-groove in this work. Thespecimens were fabricated with Ti6Al4V material at one end and Ti6Al4V/10wt%TiC at theother. The joint designs define the transition from one material to the other. The butt joint wasdesigned for an abrupt transition from one material to the other. Interlock joints were fabricatedwith interlocking strips of materials of 1mm thickness. They were made by depositing fourlayers of materials for each strip. Random joints were designed with random lengths ofinterlocking strips of 0.5mm. They were made by depositing two layers of materials for eachstrip. The scarf joint was designed with a lap angle of 45o while the included angle of the v-groove joint design is 90o.

They were made by depositing two layers of materials for eachstrip. The scarf joint was designed with a lap angle of 45o while the included angle of the v-groove joint design is 90o. The v-groove specimens were fabricated by first depositing the twoprincipal materials completely, and the groove subsequently filled up with the transition material.In this work, the groove filler material was Ti6Al4V, one of the two principal materials. Thegradient transition joint was designed to vary TiC composition linearly from zero to 10wt%TiCin Ti6Al4V at a step interval of 0.25mm over a total length of 5mm.Material-specific STL CAD files were developed for the fabrication of the dual-materialtest specimens. Two different material mixtures were automatically deposited side-by-side foreach layer following the spatial relationship specified in the machine code, which was dependenton the joint design.

Two different material mixtures were automatically deposited side-by-side foreach layer following the spatial relationship specified in the machine code, which was dependenton the joint design. The gradient transition joint specimens were fabricated with orientationsparallel to the deposition z-axis direction, as more than three axes would be required for itsdeposition in any other direction. Specimens with other joint designs were deposited withorientations perpendicular to the deposition nozzles. Single material specimens were fabricatedto determine the as-deposited tensile properties of the two base material mixtures (Ti6Al4V andTi6Al4V/10wt%TiC).

Single material specimens were fabricatedto determine the as-deposited tensile properties of the two base material mixtures (Ti6Al4V andTi6Al4V/10wt%TiC). A combination of computer numerical control (CNC) milling and wireelectrical discharge machining (EDM) were used to finish up the specimens to final size and toremove them from the substrate.(a) Butt joint(b) Gradient joint(c) Interlocking joint(d) Random interlock joint(e)Scarf joint(f) V-roove jointFigure 4: Multi-material interface designs for LENS depositionAll tensile specimens were made in accordance with ASTM E 8_E 8M. 3-point bendspecimens were sized based on ASTM C1341-06.

3-point bendspecimens were sized based on ASTM C1341-06. The flexural strengths were calculated using223bdFLf(ii)where,fflexural strengthF = load at failureL = specimen support spanb = specimen width, andd = specimen thickness.An ASTM D 2344 short beam 3-point test fixture was used for the flexural strength tests.A 50kN Tinius Olsen tensile testing machine was used for all tests at a crosshead speed of0.5mm/minute.Dual-material minimum weight structures designed based upon maximum strength andmaximum stiffness criteria were fabricated using selected material transition joint designs. Thisselection was based on the results of the joint design characterizations just described. Thestructures were meant to test the performance of the joint designs in practical applications. Theywere fabricated using Ti6Al4V as the compression member material and Ti6Al4V/10wt%TiC asthe tension member material.

Theywere fabricated using Ti6Al4V as the compression member material and Ti6Al4V/10wt%TiC asthe tension member material. Depending on the application, the compression and tensionmember materials as specified can be maintained or reversed. The Ti6Al4V/10wt%TiC wasselected for the tension members in anticipation of higher tensile strength than Ti6Al4V.Members subjected to tension fails faster than compression members under loading conditions.The use of the composite material in the tension members was intended to increase load carryingcapability before failure of the structure. An initial attempt to fabricate the structures withTi6Al4V/10wt%TiC composite as the tension member using the T- and V-shaped materialmodels shown in Fig. 5 was unsuccessful as cracks developed during deposition, mostly after theeighth layer. The cracks initiated and propagated at the Ti6Al4V/10wt%TiC composite side atthe joints. Butt and scarf joints shown in Fig.

The cracks initiated and propagated at the Ti6Al4V/10wt%TiC composite side atthe joints. Butt and scarf joints shown in Fig. 5 were used during those trials. The labels A and Bindicate mating surfaces of the T- and V- geometries during deposition to form the final shape.The cracks developed irrespective of the joint design used. However, structures fabricated with5wt%TiC composition in the tension material member did not crack.The problem with cracks necessitated a change in the design of the material models toallow for a combination of butt and interlock joints at all the material intersections as shown inFig. 6. Figure 6a is a fork shaped Ti6Al4V material model with arms extending through thetriangular shaped structure. The two lower arms in Fig. 6 are of 0.5mm thickness, while thetopmost arm is of 1mm thickness. The arms were intended to separate Ti6Al4V/10wt%TiCcomposite tension members (with material model shown in Fig.

6 are of 0.5mm thickness, while thetopmost arm is of 1mm thickness. The arms were intended to separate Ti6Al4V/10wt%TiCcomposite tension members (with material model shown in Fig. 6b) into three discrete partitionswith Ti6Al4V as 0.5mm thick inter-layers. However, during the fabrication, rather than deposittwo layers of Ti6Al4V materials consecutively at the inter-layer, they were alternated withTi6Al4V/10wt%TiC composite layers. This deposition method significantly reduces theoccurrence of cracks in the deposits. Fabricated minimum weight structures were tested with anASTM D 2344 short beam test fixture as illustrated in Fig. 7. The load at failure was obtained foranalysis.

Fabricated minimum weight structures were tested with anASTM D 2344 short beam test fixture as illustrated in Fig. 7. The load at failure was obtained foranalysis. Also, the mode of failure, especially the fracture location, was studied.Figure 5: T- and V-shaped CAD material modelsFigure 6: Material specific CAD modelsFigure 7: Structure under test using a 3-point bend test fixture6.3Results and Discussion3.1 MicrostructuresSome of the micrographs of deposited specimens are shown in Fig. 8 below. Figure 8ashows a micrograph of Ti6Al4V/10wt%TiC deposit. Figure 8b and 8c shows the butt jointinterface with some level of interlock. The interlock is due to the fact that the two materialmodels used for the dual material specimens share the same contour boundary line, and in everylayer, each of the materials is deposited at the common boundary resulting in small amount ofinterlock. A gradual transition from one material to the other can be observed in Fig. 8c.

A gradual transition from one material to the other can be observed in Fig. 8c. Figures8d and 6.8e shows sandwiched strips of Ti6Al4V/10wt%TiC composite and Ti6Al4V materialsin an interclock material joint. The micrograph of a scarf joint is shown in Fig. 8f.(a): Ti6Al4V/10wt%TiC composite showing some un-melted particles(b): Butt joint showing small interlock at the boundary contour deposit(c): Butt joint showing material mixing at the interface(d): Interlocked joint showing a sandwiched Ti6Al4V/10wt%TiC strip(e): Interlocked joint showing a sandwiched Ti6Al4V strip(f): Scarf joint showing the interface between Ti6Al4V/10wt%TiC and Ti6Al4VFigure 8: Micrographs of some of the dual-material test specimens3.2Flexural and Tensile StrengthsThe flexural strength data obtained are shown in Table 2 and graphically in Fig. 9. Theflexural strengths of the base materials (Ti6Al4V and Ti6Al4V/10wt%TiC) are included.

9. Theflexural strengths of the base materials (Ti6Al4V and Ti6Al4V/10wt%TiC) are included. It canbe seen from the data that under the bending load condition, the scarf joint design yielded thehighest average flexural strength. It also has one of the lowest strength variations. The butt jointdesign yielded the lowest average strength. The result of a statistical analysis of the data usingSAS 9.1 show that joint design (with a p-value of 0.2268) does not have statistically significanteffect on the flexural strength of the LENS fabricated specimens. It means for an applicationrequiring lateral loading, any of the inter-material joint designs can be used. The ease offabrication will therefore be a major consideration for such applications.The tensile strengths data for corresponding joints designs are shown in Table 3 and Fig.10.

The ease offabrication will therefore be a major consideration for such applications.The tensile strengths data for corresponding joints designs are shown in Table 3 and Fig.10. The results of the analysis of the data show that joint design has a statistically significanteffect on the tensile strengths of LENS fabricated dual-material structures with a low p-value of0.0002 as shown in the analysis of variance (ANOVA) Table 4a. Single material specimens withTi6Al4V and Ti6Al4V/10wt%TiC expectedly have the highest average tensile strength values.According to the Ryan-Einot-Gabriel-Welsch (REGWQ) post hoc means analysis table shown inTable 4b, both interlock and butt joint designs have statistically significantly higher averagetensile strengths than random and v-groove designs. Although the former pair yielded higheraverage tensile strengths than gradient and scarf joints, the differences are not statisticallysignificant.

Although the former pair yielded higheraverage tensile strengths than gradient and scarf joints, the differences are not statisticallysignificant. It means any of those four designs can be used in place of another in LENSfabricated dual-material structures. One of the major defects that might have resulted in lowtensile strengths recorded for the v-groove joint design is shown in Fig. 11. In this design, thetwo principal materials are first deposited completely before the transition joint material isdeposited in the groove. With the as-deposited rough surfaces of the principal materials,sometimes the laser does not have all the surfaces exposed for re-melting and deposition with thetransition material. Voids are therefore created in the process.

12. The jointlocations are marked with letters A and B. Data obtained from the loading tests are shown inTables 6 to 8. For ease of representation, the structures are labeled as follows: the maximumstrength criterion structures with 5wt%TiC in the tension members are denoted STR5; themaximum stiffness criterion structures of the same composition are STF5; and the correspondingstructures with 10wt%TiC are STR10 and STF10. Table 5 shows the member (oa, oc and acillustrated in Fig.

3) applied at the time of failure for each structure isshown in Table 6. Table 7 shows the stress on the tension members at failure, while Table 8shows the calculated strain energy density for each structure. The load applied at failure for bothmaterial composition structures, seen graphically in Fig. 13, are close in value, although they arenot directly comparable since they are of different thicknesses. The stresses on the tensionmembers (ac or ab in Fig. 3) for each structure are shown in Table 7 and Fig. 13. STR10 andSTF10 structures failed at slightly higher stresses in the tension members than correspondingSTR5 and STF5 structures. The stress values are about 50% of the tensile strength of therespective materials. The tensile strengths as experimentally determined are shown in Table 9.

The tensile strengths as experimentally determined are shown in Table 9. Itis noteworthy that no failure occurred at the joints in any of the structures irrespective of the TiCcomposition used.Table 9: Average Tensile properties of materials used MaterialYield strengthTensile strengthTi6Al4V(MPa) 1072(MPa) 1169Modulus (GPa) 111Ti6Al4V/5wt%TiC9851099114Ti6Al4V/10wt%TiC10891196154Strain energy densities shown in Table 8 were calculated based on the structure membersizes, material properties and strain values usingVvU)21((iii)Where, U = strain energy density of a structure at failure loadstress in each member at the failure loadstrain of each member at failure loadvvolume of each structure memberVtotal volume of the structure being analyzedThe stress σ in each member at the point of failure was calculated by normalizing theresolved load (based on Table 1 relationships) with respect to cross-sectional area.

With thestress obtained, the strain ε was calculated using the stiffness value for each structure member.The strain energy density data shows that the structures designed based on maximumstiffness criterion generally yielded higher average strain energy density than those designedbased on maximum strength criterion for both tension member material compositions (5wt% TiCand 10wt%TiC). This result is opposed to the results obtained in earlier work [20] in whichultrasonically consolidated structures based on maximum strength criterion yielded significantlyhigher average strain energy density than those designed based on maximum stiffness criterion.Most of the structures in the referenced work failed at the foil edge-to-edge joints at the tensionmembers as opposed to failures at the flanges for most of the LENS fabricated structures. It wasfor only one case that the maximum strength criterion structure failure occurred at one of thetension members, as shown in Fig.

It wasfor only one case that the maximum strength criterion structure failure occurred at one of thetension members, as shown in Fig. 14a. In all other cases, both for maximum strength andmaximum stiffness criteria, failures occurred at the flanges as shown in Fig. 14b.(a): Failure at a tension member on an STR10 structure(b): Failure at the flange obtained in most structuresFigure 14: Fracture locations in LENS fabricated minimum weight structures+4ConclusionsIt has been shown experimentally in this work that transition joint design does not havesignificant effect on the flexural strengths of LENS fabricated dual-material structures made ofcompatible materials. In contrast, joint design has significant effect on their tensile strengths. V-grooved and randomly interlocked joint designs yield poor tensile strengths when compared tointerlocked, butt, gradient and scarf joints.

V-grooved and randomly interlocked joint designs yield poor tensile strengths when compared tointerlocked, butt, gradient and scarf joints. Among the later four design types, the interlockdesign yielded the best average tensile strength. It also performed well under flexural loading.However, any of the four designs can be used for structural applications. None of the LENSfabricated minimum weight structures fail at the material transition joints under 3-point loadingconditions. This work has shown that several different types of joint designs may work reliablyfor multi-material components.References1Cohen D.L., Malone E., Lipson H.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDetermining the Complex Young’s Modulus of Polymer Materials Fabricated with MicrostereolithographyC. Morris*, J. M. Cormack*†, M. F. Hamilton*†, M. R. Haberman*†, C. C. Seepersad*Department of Mechanical Engineering, The University of Texas at Austin, Austin, TX 78712 †Applied Research Laboratories, The University of Texas at Austin, Austin, TX 78758AbstractMicrostereolithography is capable of producing millimeter-scale polymer parts having micron-scale features. Material properties of the cured polymers can vary depending on build parameters such as exposure time and laser power. Current techniques for determining the material properties of these polymers are limited to static measurements via micro/nanoindentation, leaving the dynamic response undetermined.

Current techniques for determining the material properties of these polymers are limited to static measurements via micro/nanoindentation, leaving the dynamic response undetermined. Frequency-dependent material parameters, such as the complex Young’s modulus, have been determined for other relaxing materials by measuring the wave speed and attenuation of an ultrasonic pulse traveling through the materials. This method is now applied to determine the frequency-dependent material parameters of polymers manufactured using microstereolithography.

This method is now applied to determine the frequency-dependent material parameters of polymers manufactured using microstereolithography. Because the ultrasonic wavelength is comparable to the part size, a model that accounts for both geometric and viscoelastic effects is used to determine the material properties using experimental data.IntroductionParts produced by additive manufacturing (AM) are increasingly utilized for applicationssuch as energy absorbing honeycomb structures, prosthetic limbs, and shock isolation systems where the response of the material to dynamic loading must be considered [1, 2, 3]. Due to the geometric design freedom introduced by AM, parts can achieve mechanical performance levels previously unattainable by other manufacturing technologies [4].

Due to the geometric design freedom introduced by AM, parts can achieve mechanical performance levels previously unattainable by other manufacturing technologies [4]. Successful prediction of the mechanical performance of parts made from AM processes requires accurate mechanical modeling which, in turn, requires precise knowledge of rate-dependent material properties of the as-built parts.The frequency dependent modulus that relates the stress developed in the material due toa dynamically applied strain is one such property. The material property describing this relationship is known as the dynamic modulus, which is frequency dependent and expressed as a complex quantity that accounts for both storage and loss of mechanical energy. The modulus of low-loss elastic materials like metals is approximately rate independent for most applications, and can therefore be described with static elastic moduli.

The modulus of low-loss elastic materials like metals is approximately rate independent for most applications, and can therefore be described with static elastic moduli. The static Young’s modulus for the uniaxial loading case is one such property that can be measured through quasi-static tensile or three point bending tests. If the material exhibits viscoelastic behavior, the mathematical description of the frequency dependent storage and loss moduli require a more generalized constitutive model [5], the parameters of which must be obtained experimentally.When a viscoelastic material is dynamically loaded, some of the imparted strain energy isstored elastically within the material while some of the energy is dissipated. The amount of426energy that is stored and dissipated can vary with the frequency of the applied load.

The amount of426energy that is stored and dissipated can vary with the frequency of the applied load. A general form of the complex modulus, 𝐸(𝜔), that captures this phenomenon is𝐸(𝜔) = 𝐸′(𝜔) + 𝑗𝐸′′(𝜔),where the real part, 𝐸′(𝜔), is the storage modulus corresponding to the frequency-dependent, elastic storage of energy, and the imaginary part 𝐸′′(𝜔) is the loss modulus that accounts for the dissipation of dynamic energy. Both the storage and loss modulus must be determined to obtain the dynamic modulus. However, the standard quasi-static test previously mentioned only captures the zero frequency component of the storage modulus. In order to obtain the complete behavior of the dynamic modulus other testing methods must be explored.One such method, ultrasonic characterization, is of particular interest to the additivemanufacturing community because it is a nondestructive way of measuring dynamic material properties over a large range of frequencies.

It is well documented that the material properties of parts produced via AM can vary across machines, parts, and even different locations of the same part [6, 7]. Therefore, a methodology must be developed that can determine material properties for individual parts both quickly and effectively; furthermore, to be applicable to a range of processes, including microstereolithography, it must be applicable to parts with small characteristic dimensions, on the order of millimeters or even smaller. In this paper, an experimental approach and analysis procedure is applied to determine the dynamic modulus of an additively manufactured part using ultrasonic characterization.

In this paper, an experimental approach and analysis procedure is applied to determine the dynamic modulus of an additively manufactured part using ultrasonic characterization. The procedure, in general, can capture the dynamic modulus for a large range of frequencies, geometries, and part sizes and was demonstrated on a rod produced by microstereolithography to determine the dynamic modulus in the ultrasonic range of 400 kHz to 1.3 MHz.MicrostereolithographyMicrostereolithography is an additive manufacturing process based on photopolymerization in which a liquid polymer solidifies when exposed to a particular wavelength of light. Figure 1 provides a schematic of a typical microstereolithography system [8]. The microstereolithography system builds parts layer by layer by activating/deactivating each pixel of a Digital Micromirror Device (DMD) so that the correct image is formed when a UV light source reflects off the device.

Then through a series of optics, the light image is greatly reduced in size and focused on the top of a build stage at the surface of a volume of liquid photopolymer. A thin layer of the photopolymer solidifies on the build plate that matches the light image. Once the layer has solidified the build stage moves downward, the liquid polymer flows over the part to form another layer, and the build process is repeated to form the final part.Figure 1: Schematic of microstereolithography system [8].427(1)The material properties of the final part produced by microstereolithography will vary based on several factors including the layer thickness, exposure time, and volume fraction of constituents [9, 10]. These parameters can vary from build to build based on desired performance. For example, some builds may require a smaller layer thickness for increased part resolution. Therefore, a metrology part should be produced for each set of build parameters to evaluate the dynamic modulus.

Therefore, a metrology part should be produced for each set of build parameters to evaluate the dynamic modulus. When developing the metrology part, the build envelope must be considered because it plays a vital role in determining what testing methods can be used. The system of interest can produce parts with resolution on the order of tens of microns with overall parts sizes of about 2.5mm x 2.5mm x 15mm. The resolution and part size can vary from system to system and the system used in this paper at the University of Akron is capable of the dimensions previously described.

The resolution and part size can vary from system to system and the system used in this paper at the University of Akron is capable of the dimensions previously described. With these geometric constraints, the testing methods to consider for determining the dynamic modulus are the Impulse Excitation Technique, Dynamic Mechanical Analysis, Nanoindentation, and ultrasonic material characterization.Dynamic Modulus DeterminationThe Impulse Excitation Technique (IET) determines the dynamic modulus by providingan impulsive load to a sample, which excites the material to vibrate at its natural frequencies [11]. The quality factor and dynamic modulus can be determined for the material by measuring the ringdown response, but only at the natural frequencies of the specimen. To obtain the modulus at other frequencies, multiple specimens must be produced, each having different geometries that permit the measurement of the dynamic modulus over a wide range of frequencies.

To obtain the modulus at other frequencies, multiple specimens must be produced, each having different geometries that permit the measurement of the dynamic modulus over a wide range of frequencies. This is rarely an efficient strategy for characterization over wide frequency ranges. Further, measurement accuracy can be negatively impacted since the material properties of additively manufactured parts often vary from part to part.The Dynamic Mechanical Analyzer (DMA) is a commercially available device that can be used to determine the dynamic modulus of viscoelastic materials [12]. The DMA provides a time-dependent load to a specimen (usually sinusoidally varying) in a temperature-controlled environment and measures the response of the part to the load. The DMA then sweeps through various ambient temperatures to obtain the frequency-dependent response of the material at those discrete temperatures.

The DMA then sweeps through various ambient temperatures to obtain the frequency-dependent response of the material at those discrete temperatures. The response for all temperature and frequency combinations can then be used to determine the dynamic modulus of the material for a wide range of frequencies and temperatures using the principle of time-temperature superposition [13]. This range of frequencies and temperatures can actually exceed the testable range if the material is thermorheologically simple, meaning regardless of the initial stress, the stress relaxation times share the same dependence on changes in temperature [14].Mixtures of photopolymers and photoinitiators have been shown to bethermorheologically complex because they exhibit multiple time-temperature shifts from multiple viscoelastic domains [15].

The principle of time-temperature superposition cannot be used for these thermorheologically complex materials so the maximum testable frequency of the DMA bounds the range of attainable information about the dynamic modulus. For commercially available DMAs this maximum testable frequency is around a few hundred hertz [16]. Furthermore, the DMA requires the specimen to be of a certain geometry to interface with its428fixtures. For certain additive manufacturing methods like microstereolithography, it may be prohibitively difficult to produce a part of adequate size to interface with the provided fixtures, therefore another testing method is needed to test these parts. It should be noted that Chartrain et al. used DMA to determine the effect of temperature on the dynamic modulus of a thin film manufactured by microstereolithography [17].

It should be noted that Chartrain et al. used DMA to determine the effect of temperature on the dynamic modulus of a thin film manufactured by microstereolithography [17]. They were able to produce a viable part because the build volume of their system was 4mm x 6mm x 35mm which allowed them to produce a larger test specimen.Another method, closely related to the DMA is the use of a nanoindentor to deform thematerial of interest. Typically, nanoindentation is used to determine the static modulus of elasticity, but if the machine is carefully calibrated and a sinusoidal indenting force is applied, the response of the material can be used to determine the frequency-dependent dynamic modulus. The careful calibration required is an extensive process and, even if carried out to ASTM standards, the results can differ when compared to the DMA [18].

The careful calibration required is an extensive process and, even if carried out to ASTM standards, the results can differ when compared to the DMA [18]. Similar to the DMA it also has a limited range of frequencies that can be used to determine the dynamic modulus limiting the efficacy of the testing.A method that allows for a large range of frequencies to be evaluated for parts of varioussizes is ultrasonic material characterization. Ultrasonic material characterization measures the propagation of an input wave pulse in a specimen and relates the response to the material properties and geometry through a forward model [19]. The forward model is the cornerstone of the method because it predicts wave propagation through a specimen. Forward models, which allow the measured response to be inverted to determine material properties, can be constructed for simple and complex geometries.

Forward models, which allow the measured response to be inverted to determine material properties, can be constructed for simple and complex geometries. Therefore, ultrasonic material characterization was selected as the method to determine the dynamic modulus of the metrology part produced using microstereolithography.Before beginning ultrasonic material characterization, a geometry must be selected forthe dynamic modulus metrology part. A natural choice for the design is a cylindrical rod due to its ability to be produced rapidly by all additive manufacturing technologies. A forward model for wave propagation in a cylindrical rod is well known and will be discussed in the next section.Ultrasonic Material CharacterizationThe simplest definition of ultrasonic material characterization is an experimental methodthat uses measurements of the speed of sound in a specimen paired with a forward model to determine the material properties of the specimen.

The forward model relates the frequency- dependent sound speed to the geometry and frequency-dependent material properties of the specimen. Accurate measurement of the sound speed paired with knowledge of specimen geometry can therefore be used to infer the material properties via a minimization of the difference between the experimental data and forward model predictions as the properties are varied [20].The simplest example of ultrasonic material characterization is the determination of afrequency-independent Young’s modulus of a material by measuring the time-of -flight, or the time it takes for a wave to travel from one point of a material to another point, in a lossless429material. The sound speed, which is more accurately called the phase speed, 𝑐p, can be determined by 𝑐p = 𝑑/𝑡 , where 𝑡 is the time-of-flight and 𝑑 is the distance traveled by the wave.

The sound speed, which is more accurately called the phase speed, 𝑐p, can be determined by 𝑐p = 𝑑/𝑡 , where 𝑡 is the time-of-flight and 𝑑 is the distance traveled by the wave. The forward model that relates the phase speed and the material properties for longitudinal pulse propagation in a thin rod is 𝑐p = √𝐸/𝜌, where E is the Young’s modulus and 𝜌 is the density. The previous equation can be inverted to obtain 𝐸 = 𝜌𝑐p be determined from measurable quantities.2 which allows the Young’s modulus toThe use of the time-of-flight method described above determines a single value of thephase speed because in that example the Young’s modulus was assumed to be frequency- independent. For most geometries and materials this is not necessarily true.

For most geometries and materials this is not necessarily true. If the Young’s modulus is frequency-dependent, the phase speed and attenuation, 𝛼, will be dependent on the frequency, 𝜔, as well [21] which results in the complex wavenumber:𝜉(𝜔) =𝜔 𝑐p(𝜔)− 𝑗𝛼(𝜔).where exp(𝑗𝜔𝑡) time dependence is assumed.The complex wavenumber is related to the material properties, frequency, and specimen geometry by a dispersion relationship. The dispersion relationship can then be inverted to relate the material properties to the phase speed and attenuation. While the dispersion relationship may be complicated and contain both real and imaginary components, the methodology for determining the material properties is nearly identical to that of the simple time-of-flight example illustrated above. First, the frequency-dependent, complex wavenumber is measured and the dispersion relationship describing the specimen is obtained.

First, the frequency-dependent, complex wavenumber is measured and the dispersion relationship describing the specimen is obtained. Next, the dispersion relationship is inverted to solve for the material properties in terms of the measured wavenumber and geometry, allowing for the determination of the dynamic modulus.Dispersion in a Solid Rod of Circular Cross-SectionFor the case of axisymmetric longitudinal wave propagation in a viscoelastic rod withcircular cross-section, the dispersion relationship is governed by two factors: (i) the viscoelastic constitutive relationship of the material, and (ii) the geometry of the rod. Pochhammer and Chree were the first to independently describe the dispersion relationship for an infinite, elastic cylindrical rod, which was later modified by Zhao and Gary to incorporate viscoelasticity [22, 23, 24].

The dispersion relationship describing axisymmetric longitudinal wave propagation in an infinite, viscoelastic, cylindrical rod is2𝐴 𝑟(𝐵2 + 𝜉2)𝐽1(𝐴𝑟)𝐽1(𝐵𝑟) − (𝐵2 − 𝜉2)2𝐽0(𝐴𝑟)𝐽1(𝐵𝑟) − 4𝜉2𝐴𝐵𝐽1(𝐴𝑟)𝐽0(𝐵𝑟) = 0,where the coefficients 𝐴 and 𝐵 are defined by𝐴2 =𝜌𝜔2 4 3𝐾 +𝜇− 𝜉2,and430(2)(3)(4)𝐵2 =𝜌𝜔2 𝜇− 𝜉2.In Eq. (3) r is the radius of the rod, 𝐽𝑛(𝑧) is the 𝑛𝑡ℎ order Bessel function of the first kind, K is the bulk modulus, and 𝜇 is the shear modulus of the material. The bulk modulus and the shear modulus are related to the Young’s modulus by 𝐸 = 9𝐾𝜇/(3𝐾 + 𝜇) [25]. Note for |𝜉𝑟| ≪ 1 and as 𝜔 → 0, Eq. (3) reduces to the simple case of longitudinal pulse propagation in a thin rod with phase speed 𝑐p = √𝐸/𝜌, as discussed in the previous section.In the derivation of Eq. (3) it is assumed that the rod is infinitely long.

(3) it is assumed that the rod is infinitely long. In practice then,the length of the rod, l, must be much larger than the wavelength, λ, whereλ =2𝜋𝑐p 𝜔≪ 𝑙.Therefore, lower frequency measurements may not be accurately described by Eq. (3), if the wavelength is on the order of the rod length. This condition must be considered when investigating rods produced by additive manufacturing systems because the build envelope determines the maximum rod length. It also should be noted that Eq. (3) cannot be easily inverted analytically as in the earlier example so that the material properties are given in terms of the frequency, wavenumber, and geometry. A numerical solver must be used to obtain the material properties once the other parameters are determined for each frequency. For a given frequency, there may be multiple solutions to Eq. (3), each corresponding to different modes of wave propagation along the axis of the rod.

For a given frequency, there may be multiple solutions to Eq. (3), each corresponding to different modes of wave propagation along the axis of the rod. The existence of multiple modes requires careful treatment when measuring the relationship between wavenumber and frequency. Figure 2 shows equivalent representations of a dispersion plot for an elastic rod where in Figure 2(a) the horizontal axis is the frequency and the vertical axis is the phase speed and for Figure 2(b) the horizontal axis is the real component of the wavenumber and the vertical axis is the frequency.Figure 2: Plot of the (a) phase speed as a function of frequency and (b) the real component of inverse wavelength as a function of frequency when multiple modes are present for r = 1.25 mm, E = 4.5 GPa, and ν = 0.375.431(5)(6)As shown in Figure 2, when a cylindrical rod is excited, multiple modes may propagatethrough the rod and appear in the complex wavenumber measurement.

The following section discusses how this measurement of the complex wavenumber is performedDetermining Complex WavenumberDetermining the complex wavenumber requires measuring both the frequency-dependent phase speed and wave attenuation. This is accomplished by measuring ultrasonic wave packets at a range of positions as they propagate along the axis of the rod to obtain spatial and temporal information about the propagating wave. Performing a two-dimensional (2D) fast Fourier transform (FFT), with one dimension being time and the other space, one obtains the complex amplitude of the propagating pulses for frequency, wavenumber pairs. The data is easily visualized with a surface plot where the horizontal axis is the real component of the wavenumber, the vertical axis is the frequency, and the magnitude is the amplitude of the transformed signal. This representation is commonly referred to as the dispersion diagram of the elastic waveguide.

This representation is commonly referred to as the dispersion diagram of the elastic waveguide. If only axisymmetric modes are present, the amplitude of the surface plot is dramatically higher along curves that satisfy the dispersion relationship Eq. (3). The degree of agreement between the values calculated with Eq. (3) (the forward model) and the experimental data depends primarily on the accuracy of the material properties provided as inputs to the forward model. As previously discussed, multiple modes may be present, so filtering of the data may be required if one wishes to isolate any specific propagating mode. Once a single mode is isolated from the dispersion diagram, the frequency-dependent phase speed for the mode of interest can be extracted by determining the real component of the wave number for each frequency where the amplitude of the signal is highest.

This allows for numerical determination of the phase speed.With the isolated mode, it is then possible to use the filtered data to determine the imaginary component of the wavenumber. An inverse Fourier transform of the filtered data can be taken to yield the amplitude of the selected mode in time and space. For each measurement separated by a distance Δ𝑥, a FFT can be taken in the time domain. Then the transfer function 𝐻(𝜔) between adjacent measurements, 𝑈̃2(𝜔, 𝑥) and 𝑈̃1(𝜔, 𝑥), is defined to be𝐻(𝜔) =𝑈̃2 𝑈̃1= 𝑒−𝛼Δ𝑥𝑒−𝑗𝜔 𝑐𝑝Δ𝑥.The magnitude of the transfer function and knowledge of distance between measurement locations can then be used to determine the attenuation coefficient using the expression𝛼 = −1 Δxln|𝐻(𝜔)|.The method described above provides a means to determine the real and imaginary components of the complex wavenumber from experimentally obtained data.

That information is then used to determine the material properties through inversion of the dispersion relationship Eq. (3).The complex wavenumber can be constructed at each frequency where attenuation and phase speed information is available through Eq. (2). The complex wavenumber and frequency432(7)(8)are substituted into the dispersion relationship of Eq. (3) which is then separated into its real and imaginary components forming a system of two simultaneous equations. The material properties are then determined by obtaining values of the bulk modulus and shear modulus that minimize the magnitude of both equations through the use of the fsolve function in MATLAB®. This methodology was applied to a rod produced via microstereolithography. The experimental setup and results are discussed in the next section.Experimental Setup and ResultsAn ultrasonic transducer, laser Doppler vibrometer (LDV), and linear stage were used toobtain spatial and temporal measurements.

Figure 3 presents a schematic of the experimental setup where a rod is placed between two ultrasonic transducers and the LDV is positioned so that the laser path is perpendicular to the axis of the rod. The rod and transducers are mounted on a linear stage that allows the sample to be moved in the axial direction. The transducer transmits a broadband ultrasonic pulse that travels axially through the rod. While the pulse propagates axially, the surface of the rod contracts and expands radially due to the Poisson effect. The radial surface velocity is measured by the LDV. Multiple measurements along the rod axis can be taken with the LDV by incrementally moving the rod and transducer setup with the linear stage.

The radial surface velocity is measured by the LDV. Multiple measurements along the rod axis can be taken with the LDV by incrementally moving the rod and transducer setup with the linear stage. This generates both spatial and temporal information of ultrasonic wave propagation along the rod.Figure 3: Schematic of experimental setup for surface velocity measurement of rod.It should be emphasized that this experimental setup has major advantages compared to other arrangements that could obtain spatio-temporal data. Two setups will briefly be discussed that could be implemented if the LDV were not available, however both methods introduce significant errors. In the absence of the LDV, the second ultrasonic transducer serves as a receiver for the transmitted signal of the first transducer. In one alternative experimental setup, spatio-temporal data could be collected by shortening/cutting the rod incrementally and measuring the received pulse for each length.

In one alternative experimental setup, spatio-temporal data could be collected by shortening/cutting the rod incrementally and measuring the received pulse for each length. Due to variation in loading conditions, the signal transmitted to the receive transducer will vary from measurement to measurement reducing the credibility of the results. A second alternative experiment uses a single rod of one length. The second transducer measures the first arrival of the wave as well as the second arrival that comes after reflection from the exciting transducer. Material properties are obtained through comparison of the two signals. The second arrival is due to an impedance mismatch between the transducer, coupling gel, and rod at the interface, therefore additional unknowns are introduced433to the measurement. Furthermore, if the material attenuates strongly, the signal from the reflected signal will be too small to detect.

Furthermore, if the material attenuates strongly, the signal from the reflected signal will be too small to detect. For viscoelastic rods, like the ones selected, the reflected signal is far too weak to be measured above the noise level.To validate the experimental and inversion approach, measurements were made with amaterial having known dynamic modulus in the ultrasonic range. Tests were therefore performed on a 12.7 mm diameter rod of acrylic (PMMA) with a length of 30 mm. The ultrasonic transducer was a 1 MHz Olympus V603-RB transducer with a nominal bandwidth of 0.69-1.29 MHz. The LDV was a Polytec OFV-505, and data was captured using a 60 MHz LeCroy WaveAce 102 digital oscilloscope. The linear stage allowed for a spatial resolution of 0.0254 mm. Spatial and temporal data were collected for the acrylic rod and the results were compared to the values previously determined in the literature.

Spatial and temporal data were collected for the acrylic rod and the results were compared to the values previously determined in the literature. The real component of the dynamic modulus for acrylic has been shown to be nearly constant in the ultrasonic range with a value of 4.9 GPa. The imaginary component of the dynamic modulus is approximately linear with frequency with a slope of 1.31 dB/cm/MHz [26, 27]. Using these values and a Poisson value of 0.37, the multimodal, real-valued dispersion curves were generated and overlaid on the results of the spectral transform of the LDV data as shown in Fig. 4.

Using these values and a Poisson value of 0.37, the multimodal, real-valued dispersion curves were generated and overlaid on the results of the spectral transform of the LDV data as shown in Fig. 4. There is excellent agreement between the experimental results and established values, especially for the first propagating mode.Figure 4: Experimental dispersion curves obtained from an acrylic rod (surface) with real components of dispersion curves determined from established values (white).Following the successful validation of the method with the acrylic rod, a rod was produced via microstereolithography with a diameter of 2.47 mm and a length of 15 mm. The polymer blend used to produce the rod was CD9021/HDDA (70w/30w), with 1 w% DMPA as the photo-initiator, and 0.15 w% Tinuvin 327® as the light absorber. After exciting the rod with the ultrasonic transducer, the radial surface velocity at various axial locations on the rod was measured with the LDV.

After exciting the rod with the ultrasonic transducer, the radial surface velocity at various axial locations on the rod was measured with the LDV. A total of 250 measurements were taken at 0.0508 mm increments and every tenth result is shown in the waterfall plot of Fig. 5. The horizontal axis is time and the vertical axis indicates distance from the first measurement taken. The waterfall plot shows that434pulse shape changes drastically as it progresses along the rod due to attenuation and dispersion effects in the waveguide.Figure 5: Waterfall plot demonstrating propagation of wave through rod.The dispersion diagram associated with the data is generated using the 2D FFT as discussed in the previous section and the results are shown in Fig. 6.Figure 6: Plot of frequency vs wavenumber obtained by 2D FFT of time-space data.To generate Fig. 6 a 250-point Hamming window was used in the spatial domain whentaking the 2D FFT of the time-space data.

6 a 250-point Hamming window was used in the spatial domain whentaking the 2D FFT of the time-space data. Figure 6 shows that significant signal is being obtained for frequencies as low as 0.15 MHz and as high as 1.6 MHz. The wavelengths at the lower frequencies are on the order of the specimen length of 15 mm so the dispersion relationship at those frequencies likely does not fit the model of Eq. (3) due to the development of standing waves. From Fig. 6 it is also be evident that multiple modes are propagating through the rod. For initial analysis of the material properties, only information regarding the first mode of propagation is desired. To isolate the first mode, the piecewise, Gaussian filter with a standard deviation of 0.15 mm-1 shown in Fig. 7 was applied to the spectral data to obtain Fig. 8.435Figure 7: Plot of filter applied to spectral data.Figure 8: Filtered plot of frequency vs wavenumber for the first mode.With the filtered plot of Fig.

8.435Figure 7: Plot of filter applied to spectral data.Figure 8: Filtered plot of frequency vs wavenumber for the first mode.With the filtered plot of Fig. 8, the real component of the wave number was obtained byselecting the wavenumber at each frequency with the highest signal amplitude. The curve depicting these results of the real component of the wavenumber against frequency is shown in Fig. 9 as well as the 5th order polynomial fit to the real dispersion curve data with an R2 value of 0.99.

9 as well as the 5th order polynomial fit to the real dispersion curve data with an R2 value of 0.99. The phase speed is also plotted in Figure 11(a) which was determined from the real part of the wavenumber using the relation 𝑐p = 𝜔\Re(𝜉).Figure 9: Real component of the wavenumber extracted from the wavenumber-frequency plot with regression curve.436The filtered spectral data was then inverted with a 2D IFFT to obtain a waterfall plot just like that of Fig 5 but with data containing information associated with the first mode contribution to the ultrasonic pulse, which is shown in Fig. 10.Figure 10: Waterfall plot of the first mode propagation.Each spatial measurement in Fig. 10 was then transformed to the temporal frequency domain and the transfer function between adjacent spatial measurements was obtained using Eq. (7). Eq. (8) was then used to obtain the frequency-dependent attenuation shown in Fig. 11(b).

(7). Eq. (8) was then used to obtain the frequency-dependent attenuation shown in Fig. 11(b). One standard deviation of variation is shown to demonstrate the low level of variance in the data.Figure 11: Plot of the (a) phase speed as a function of frequency and (b) attenuation as afunction of frequency.437The attenuation of the material is characteristic of a viscoelastic material where the attenuation, in general, increases with frequency. With the real and imaginary components of the wavenumber determined, the dynamic modulus was determined by inverting the dispersion relation Eq. (3) and assuming a Poisson ratio of 0.375. Results are presented in Fig.

(3) and assuming a Poisson ratio of 0.375. Results are presented in Fig. 12 where the variation in the dynamic modulus, plotted in green, is derived from the variation in the attenuation and the blue line is determined with the mean attenuation values.Figure 12: Plot of the (a) real component of dynamic modulus and (b) imaginary component of dynamic modulus.The dynamic modulus was obtained for the frequency range of 0.4 MHz to 1.3 MHzalthough there was information outside this range. This range was chosen to match the nominal bandwidth of the transducer as well as to ensure the assumptions regarding the dispersion relationship were not violated.

This range was chosen to match the nominal bandwidth of the transducer as well as to ensure the assumptions regarding the dispersion relationship were not violated. The values of the moduli are well within reasonable limits as the static modulus for both HDDA and CD9021 have been both experimentally determined to be around 1 GPa [28, 29].To corroborate the value of the storage modulus of 0.59 MPa at 400 kHz that was determined through ultrasonic material characterization, a secondary experiment was designed to measure the storage modulus at lower frequencies.

The experimental setup is shown in Figure 13438where the microstereolithography rod was mounted as a cantilever beam and excited by a shaker table to measure the resonance frequency.Figure 13: Experimental setup showing microstereolithography rod mounted as a cantilever to measure the storage modulus at lower frequencies.The same LDV used previously measured the velocity at the base of the cantilever beamas well as the tip of the cantilever beam while a logarithmically increasing frequency sweep from 1 Hz to 4500 Hz was performed over 120 s. By computing the transfer function between the top velocity and base velocity, shown in Figure 14, the natural frequency of the rod, 𝑓n, was found to be 1.63 kHz.Figure 14: Transfer function between the base velocity (measurement location 1 in Fig. 13) and tip velocity (measurement location 2 in Fig.

13) and tip velocity (measurement location 2 in Fig. 13).The storage modulus at the natural frequency was determined to be 0.56 MPa from 2𝜋𝑓n 3.522 𝐴𝜌𝑙4 𝐼𝐸′ = (),where the length of the rod, l, was 12 mm as opposed to the original 15 mm due to the size of the clamps, A is the cross-sectional area, and I is the area moment of inertia [30]. The experimentally obtained storage modulus at the resonance frequency of this rod (1.63 kHz) is well within the range of variance for the storage modulus at 400 kHz. The agreement in values indicates that the ultrasonic material characterization method is accurately predicting the storage modulus though caution should be used due to the difference in frequency of the corroborating results.439(9)ConclusionA methodology for determining the dynamic modulus of parts produced by additive manufacturing was applied to a rod produced by microstereolithography.

By modeling the dispersion relationship for the part of interest and then measuring the frequency-dependent attenuation and sound speed experimentally, the dynamic modulus was determined. For the microstereolithography rod, the dynamic modulus was obtained in the ultrasonic range of 0.4 MHz -1.3 MHz by utilizing the dispersion relationship for an infinite cylindrical rod. This range of frequencies is much larger than that of most methods to determine the dynamic modulus and quasi-static modulus like DMA, IET, and nanoindentation.Future work will focus on obtaining the dynamic modulus at higher and lower frequencies than those already determined. Determination of the modulus at lower frequencies while simultaneously obtaining a wide frequency range of dynamic moduli can be accomplished by using a longer rod length or different dispersion relationship.

To increase the upper limit on the frequency range, information regarding higher mode wave propagation can be explored.AcknowledgementsThe authors would like to acknowledge support from the National Science Foundation under Grant No. CMMI -1435548. We would also like to acknowledge Sumanth Kashyap and Dr. Jae-Won Choi from the University of Akron and Dr. Eric MacDonald from Youngstown State for their input and production of the microstereolithography rods. Any opinions, findings and conclusions or recommendations expressed in this material are those of the authors and do not necessarily reflect the views of the sponsor.440References[1] D. M. Correa, T. Klatt, S. Cortes, M. R. Haberman, D. Kovar and C. C. Seepersad,"Negative Stiffness Honeycombs for Recoverable Shock Isolation," Rapid Prototyping Journal, vol. 21, no. 2, pp. 193-200, 2015.[2] N. P. Fey, G. K. Klute and R. R.

Kovar and C. C. Seepersad,"Negative Stiffness Honeycombs for Recoverable Shock Isolation," Rapid Prototyping Journal, vol. 21, no. 2, pp. 193-200, 2015.[2] N. P. Fey, G. K. Klute and R. R. Neptune, "The Influence of Energy Storage and ReturnFoot Stiffness on Walking Mechanics and Muscle activity in Below-Knee Amputees," Clinical Biomechanics, vol. 26, no. 10, pp. 1025-1032, 2011.[3] B. A. Fulcher, D. W. Shahan, M. R. Haberman, C. C. Seepersad and P. S. Wilson,"Analytical and Experimental Investigation of Buckled Beams as Negative Stiffness Elements for Passive Vibration and Shock Isolation Systems," Journal of Vibration and Acoustics, vol. 136, no. 3, p. 031009, 2014.[4] B. Derby, "Printing and Prototyping of Tissues and Scaffolds," Science, vol. 338, no. 6109,pp. 921-926, 2012.[5] H. T. Banks, S. Hu and Z. R. Kenz, "A Brief Review of Elasticity and Viscoelasticity for Solids," Advances in Applied Mathematics and Mechanics, vol. 3, no. 1, pp. 1-51, 2011.[6] D. L.

G. Herbert, W. C. Oliver and G. M. Pharr, "Nanoindentation and the DynamicCharacterization of Viscoelastic Solids," Journal of Phyics D: Applied Physics, vol. 41, no. 7, p. 074021, 2008.[19] F. A. Firestone, "Flaw Detecting Device and Measuring Instrument". US Patent 2,280,226,21 April 1942.[20] D. E. Chimenti and A. H. Nayfeh, "Ultrasonic Reflection and Guided Wave Propagation in Biaxially Laminated Composite Plates," The Journal of the Acoustical Society of America, vol. 87, no. 4, pp. 1409-1415, 1990.[21] M. O'Donnell, E. T. Jaynes and J. G. Miller, "Kramers-Kronig Relationship Between Ultrasonic Attenuation and Phase Velocity," The Journal of the Acoustical Society of America, vol. 69, no. 3, pp. 696-701, 1981.[22] C. Chree, "The Equations of an Isotropic Elastic Solid in Polar and Cylindrical Co-ordinates their Solution and Application," Transactions of the Cambridge Philosophical Society, vol. 14, p. 250, 1889.[23] L.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDeveloping Additively Manufactured Iron Powder-filled PLA CompositesJeevan Kumar*, Orkhan Huseynov*, Ismail Fidan¥, and Mustafa Rajabali#Department of Mechanical Engineering, Tennessee Technological University, Cookeville, TN 38505 ¥Department of Manufacturing and Engineering Technology, Tennessee Technological University, Cookeville, TN 38505 #Department of Physics, Tennessee Technological University, Cookeville, TN 38505 AbstractThe Low-Cost Metal Material Extrusion (LCMMEX) process has gained attention in recent years to produce metallic parts with complex designs. High Iron concentration composite materials have been extensively researched to improve their properties and functionality for advanced manufacturing applications.

High Iron concentration composite materials have been extensively researched to improve their properties and functionality for advanced manufacturing applications. This study aims to develop knowledge blocks for producing composite parts with high iron percentages and analyze their physical properties based on print parameters. A rectangular sample is manufactured using the Material Extrusion (MEX) process, with variations in layer height, infill density, and print speed. The investigation shows that a number of parameters affects the change in surface roughness, weight, and dimensional accuracy of the printed parts. Furthermore, an increase in the infill percentage leads to a significant increase in magnetic flux.

Furthermore, an increase in the infill percentage leads to a significant increase in magnetic flux. This research study provides insights into the influence of print parameters on the properties of high-iron-filled composite parts, ideal for high-density applications.IntroductionAdditive Manufacturing (AM) is a technology that enables the production of three- dimensional objects by adding material layer by layer. Its capabilities in creating complex geometries and customized parts with high accuracy and precision make it valuable across industries such as aerospace, automotive, medical, and consumer goods [1].The American Society for Testing and Materials (ASTM) has standardized seven cores AM processes that utilize the layer-by-layer approach to shape desired geometries and produce final parts as per customer requirements. While the materials used may vary across these processes, the fundamental fabrication methods remain consistent.

While the materials used may vary across these processes, the fundamental fabrication methods remain consistent. For a comprehensive understanding of the layer-by-layer AM process, Figure 1 offers a detailed visual representation with corresponding descriptions, serving as a helpful reference [2].Among the seven common AM technologies, which are presented in Figure 2. MEX is widely applicable in various industries today [3]. It is a popular method for producing lightweight parts with customized shapes. Additionally, MEX is ideal for rapid prototyping and fulfilling diverse customer needs [4]. MEX employs various materials like Acrylonitrile butadiene styrene (ABS), Polylactic acid (PLA), nylon, glass-filled nylon, High Impact Polystyrene (HIPS), and Polyvinyl Alcohol (PVA) [5, 6]. The process involves a machine that extrudes filament from a spool through a nozzle, using a stepper motor to place the material layer by layer onto the build plate.

The process involves a machine that extrudes filament from a spool through a nozzle, using a stepper motor to place the material layer by layer onto the build plate. The layer thickness is in the range of 0.07 mm to 0.8 mm depending203on the part built and its settings. The final part is built in the x, y, and z directions, and once printing is complete, the machine produces a finished product depending on the requirements [7].Figure 1 MEX Printer LayoutFigure 2 Seven Different Types of AMAM has undergone significant advancements since its inception in the 1980s. Figure 3 depicts the timeline of AM technology, highlighting significant milestones and innovations in the field [8–10].204Figure 3 Chronological Representation of AM AdvancementsIn 1984, Chuck Hull invented stereolithography (SLA), the first commercial AM process. In 1987, Scott Crump developed Fused Deposition Modeling, a widely used AM technology.

In 1987, Scott Crump developed Fused Deposition Modeling, a widely used AM technology. In the 1990s, selective laser sintering (SLS) and laminated object manufacturing (LOM) were invented, offering new possibilities in AM technology. The first commercial AM systems were also introduced in this decade, expanding the technology's reach. In the 2000s, advancements in AM technologies were made through improved software, materials, and equipment. Direct metal laser sintering (DMLS) and electron beam melting (EBM) were invented, expanding the range of materials available for AM. In recent years, AM technology has significantly improved speed, accuracy, and affordability. In 2010, HP developed Multi Jet Fusion, which enables the production of functional parts at a fraction of the cost and time of traditional manufacturing methods.

In 2010, HP developed Multi Jet Fusion, which enables the production of functional parts at a fraction of the cost and time of traditional manufacturing methods. Carbon 3D company also introduced Digital Light Synthesis (DLS) in 2015, which uses light and oxygen to produce parts with mechanical properties like injection-molded plastics. As technology improves widely, we are using desktop printers for convenience [11]. Today, AM continues to evolve, with advancements in materials, software, and hardware driving its growth. Technology has the potential to revolutionize manufacturing by offering unprecedented design freedom, reduced lead times, and cost savings [9, 11].Several studies have explored the influence of various factors on the magnetic and mechanical properties of additively manufactured parts using different materials and technologies. Buchanan et al. and Ralchev et al.

Buchanan et al. and Ralchev et al. both investigated the impact of infill density and layer height on energy consumption, mass, magnetic flux density, and print time. They found that adjusting these parameters can lead to reduced print time, energy consumption, and material usage [12, 13].In a similar vein, Bollig et al. examined the effect of the internal fill factor on magnetization and observed that higher fill factors result in decreased magnetization. Patton et al. explored the relationship between infill orientation, macroscopic shape, and magnetic susceptibility, discovering that infill orientation has a significant impact on etic properties [14].205Furthermore, Watson et al. and Henderson et al. studied the influence of print parameters and deposition direction on the magnetic properties of printed parts. Watson et al. observed that changing print parameters affected the magnetic field response, indicating the importance of filament orientation and processing conditions.

Watson et al. observed that changing print parameters affected the magnetic field response, indicating the importance of filament orientation and processing conditions. Henderson et al. found that printing in the presence of a magnetic field enhanced the magnetic properties of the samples compared to those printed without a magnetic field [15, 16].Guan et al. investigated composites filled with carbonyl iron particles, demonstrating that saturation magnetization does not have a linear relationship with filler content. They also observed improved mechanical strength in the composites compared to traditional methods [17].Mohan et al. focused on the mechanical properties of MEX-printed parts and found that infill pattern and build orientation significantly affect mechanical performance, with denser infill patterns and vertical orientations resulting in superior mechanical properties [9].Laureto et al.

investigated copper fill, bronze fill, magnetic iron PLA, and stainless- steel PLA composites, reporting significant improvements in thermal conductivity compared to traditional materials. Oksiuta et al. studied PLA composites modified with Mg, Fe, and Polyethylene, observing increased hardness and tensile strength in iron-based specimens. Oyelaja et al. focused on PLA composites with carbon and iron particles, enhancing electrical conductivity and magnetic properties, while developing a neural network model for accurate prediction of composite properties [18–20].Yongquan et al. explored carbonyl iron powders, finding increased density, hardness, and bending strength with higher sintering temperature and pressure. Dayue Jiang et al. investigated PLA/iron composite scaffolds for bone tissue engineering, demonstrating improved mechanical properties and degradation resistance. Huseynov et al.

explored the fabrication of electrical motor parts using fused filament fabrication with an 80% iron ore composite. They assessed the processability and thermal properties of the filaments and established a two-step binding and sintering process, which resulted in a minimal loss of polymer decomposition and shrinkage [8].According to the reviewed literature, the thermophysical and geometric properties of 3D Printed PLA parts with a high level of iron concentration have not received extensive research attention, and the influence of post-processing on their properties remains largely unexplored.206The work presented here seeks to understand how different combinations of process parameters affect these properties, providing insights into optimizing the LCMMEX process for enhanced performance and quality. This unique study will contribute to the AM body of knowledge by adding new findings to produce low-cost iron parts in a few different ways [14].

This unique study will contribute to the AM body of knowledge by adding new findings to produce low-cost iron parts in a few different ways [14]. Additionally, this study aims to analyze the physical characterization of an Iron-PLA composite material[24].Materials and MethodologyIn pursuit of our research goal, this research is designed as a structured experiment with twelve distinct parameters, as outlined in Table 1. For specimen printing, this research has maintained precise conditions: nozzle temperature at 225°C, bed temperature at 65°C, and a nozzle diameter of 0.8mm. The specimens are consistently sized at 50mm x 25mm x 3mm. The material used for printing is a High percentage of Iron-PLA. Which is commercially available on the market from a virtual foundry website [25]. The parameters that have been varied include the infill percentage, layer height, and print speed.

Which is commercially available on the market from a virtual foundry website [25]. The parameters that have been varied include the infill percentage, layer height, and print speed. All the samples have been printed under atmospheric room temperature conditions.The MEX process involves several steps, including the preparation of the 3D Computer-Aided Design (CAD) model, slicing of the model into individual layers, and converting the data into G-code, which controls the movement of the extruder a schematic diagram of generating G-code is shown in Figure 4. The extruder heats and melts the thermoplastic material, which is then deposited onto the building platform.

The extruder heats and melts the thermoplastic material, which is then deposited onto the building platform. Once a layer is completed, the build platform is lowered, and the process is repeated until the desired object is produced.Figure 4 Steps of the Printing ProcessTable 1 DOE TableInfill Density (%)Layer Height (mm)Print Speed (mm/s)1000.1451000.2451000.1601000.260600.145600.245600.160207600.260200.145200.245200.160200.260Various equipment’s were used to record the output variables and their respective responses in the study. The printing time variable was documented in hours, representing the duration of the printing process from the initial layer to the completion of the fabricated part. To ensure accuracy and avoid any influence from the initial print temperature, the print time was measured from the point when material extrusion commenced.

To ensure accuracy and avoid any influence from the initial print temperature, the print time was measured from the point when material extrusion commenced. This timing adjustment was necessary because it takes less time for a printer to heat a hot nozzle from 200°C to 220°C compared to heating a nozzle from 20°C to 220°C [26, 27].The mass of each print was individually measured using a Mettler Toledo PL602 S precision balance. Before each measurement, the scale was zeroed (tared) to eliminate any measurement errors. Dimensional measurements, including length, width, and height, were collected using Ingress protection (IP)54 digital calipers on the printed parts. Height measurements were taken along the z-axis of the print direction, width measurements along the x-axis, and length measurements along the y-axis.

Height measurements were taken along the z-axis of the print direction, width measurements along the x-axis, and length measurements along the y-axis. These measurements were then compared to the designed dimensional parameters to determine the percentage difference between the measured values and the process parameters [8]Surface roughness, which indicates changes in the surface texture, was measured in micrometers using the Mitutoyo SJ-210 Surface Roughness Equipment. Specifically, the Ra surface roughness was measured, which represents the average deviation of a measured surface. The samples were measured by keeping the part stationary while the sensor traced over the surface which are in 45° orientation.

The samples were measured by keeping the part stationary while the sensor traced over the surface which are in 45° orientation. Measurements were taken on the top of the print relative to the print bed, as well as the width and length of the print [7, 28, 29].Magnetic flux density, which characterizes the change in the total magnetic field passing through a specific surface area, was measured using the FW Bell Model 640 Gaussmeter. This device measured the magnetic flux density of the samples. A sample was placed perpendicularly between the magnet and the three-axis probe, and the resulting change in the magnetic field was measured to determine the magnetic flux density which are in different orientations. The three-axis probe was set to x, y, and z-direction, aligning with the length of the probe.

The three-axis probe was set to x, y, and z-direction, aligning with the length of the probe. The probe holder ensured the probe remained stationary, while the magnet and test sample were kept at a consistent distance of 7.5 mm from each other [9, 12, 16, 17, 30].Thermal conductivity is a material property that measures its ability to conduct heat. It describes how quickly heat can flow through a material when there is a temperature difference. It is typically measured in watts per meter-kelvin (W/mK). High thermal conductivity allows for fast heat transfer, while low thermal conductivity indicates poor heat conduction. In AM, thermal conductivity is crucial for managing heat during the printing process, selecting208appropriate materials, optimizing energy consumption, and ensuring the performance and functionality of printed parts.

The heat transfer rate in one-dimensional conduction through a cylindrical cross-section is given by Fourier's law, where the thermal conductivity and temperature gradient play key roles. Experimental setups involve thermocouples, brass cylinders, and water cooling to measure thermal conductivity accurately [18, 19, 27, 31].To conduct the ANOVA analysis, the relevant R code was written and uploaded into the software. The code included specifications for the response variable and input variables of interest. As part of the analysis, any outliers present in the dataset were identified and subsequently removed from the analysis data. This step was taken to ensure the robustness and accuracy of the results obtained from the ANOVA analysis.Steps for analysis of ANOVA in R software: 1. Install and Load Required Packages: Begin by installing and loading the necessary R packages for ANOVA analysis, such as stats and car.

Install and Load Required Packages: Begin by installing and loading the necessary R packages for ANOVA analysis, such as stats and car. Use the install Packages () and library () functions to install and load the packages, respectively.2. Import and Prepare Data: Import the dataset containing the relevant variables for analysis, including the response variable (e.g., surface roughness, weight, dimensional accuracy) and the factors of interest (e.g., infill density, layer height). Ensure that the data is properly formatted and stored in a suitable structure, such as a data frame.3. Check Assumptions: Assess the assumptions of ANOVA, such as normality and equal variance. Use diagnostic plots, such as residual plots (plot () function) or normality plots (qqPlot () function from the car package), to evaluate these assumptions. If the assumptions are violated, consider transformations or alternative non-parametric tests.4.

If the assumptions are violated, consider transformations or alternative non-parametric tests.4. Perform ANOVA Analysis: Use the aov () function in R to conduct the ANOVA analysis. Specify the formula using the ~ symbol, with the response variable followed by the factors of interest. For example, aov_model <- aov(response variable ~ factor1 + factor2, data = your_data).5. Check ANOVA Results: Utilize the summary () function to examine the ANOVA results. This will provide an overview of the analysis, including the F-statistic, p- value, and degrees of freedom for each factor. Assess the significance of each factor's effect on the response variable based on the p-values.6. Perform Posthoc Tests: If the ANOVA results indicate significant differences between factor levels, conduct posthoc tests to determine specific pairwise differences. Commonly used posthoc tests in R include the TukeyHSD () function from the stats package or the glht () function from the multicopy package.

Commonly used posthoc tests in R include the TukeyHSD () function from the stats package or the glht () function from the multicopy package. These tests allow for pairwise comparisons and provide adjusted p-values. Results and DiscussionThe subsequent sections will focus on the analysis of process-related parameters to provide practical insights for practitioners in the field of AM. These findings aim to assist AM professionals by offering valuable guidance based on statistical analysis and data-driven results. The analysis was performed using the ANOVA method with the aid of R software. By following this approach, the analysis aimed to uncover significant relationships and variations among the process-related parameters in AM.

By following this approach, the analysis aimed to uncover significant relationships and variations among the process-related parameters in AM. The findings obtained from the ANOVA analysis, conducted using R software, will contribute to the understanding and improvement of AM209processes, enabling practitioners to make informed decisions based on the statistical evidence derived from the data.Analysis of Magnetic SusceptibilityTable 2 reveals the significant impact on magnetic susceptibility due to all examined input variables. Infill density demonstrates the greatest influence, with 81.3% of the observed variation as shown in Figure 5. Figure 6 and Figure 7 display the layer height and print speed contribution, which is less than 5% each to the overall variation. Higher infill density leads to increased magnetic susceptibility due to a greater amount of iron material in the print.

The most influential factor determining thermal conductivity is infill density. High iron samples exhibit a thermal conductivity of 10.5 W/mK, with a low mean square error for infill density is 12.48 and a p-value of 0.0001882, indicating a highly significant difference. The data variability is illustrated in Figure 9 and Figure 10. Although print speed and layer height also contribute, infill density has a greater impact on thermal conductivity which was shown in Figure 8 it has an R-squared value of 0.834.

However, print speed and layer height do not demonstrate a significant effect in this regard. Specifically, a higher infill density is associated with an increase in mass, and infill density alone accounts for 75.35% of the observed variability, as shown in Figure 11.On the other hand, the influence of layer height and print speed on mass is found to be negligible, explaining less than 1% of the variability, as illustrated in Figure 12. Interestingly, increasing layer height tends to decrease the variability in mass, which could be attributed to factors such as accidental over-extrusion or reduced voids between layers.

The analysis indicates that layer height exhibits a significant impact on print time, as evidenced by its substantial sum of squares, high F value (689.7959), and exceptionally low p-value (4.746e-09). These findings suggest that variations in layer height significantly affect the time required for printing. On the other hand, infill density demonstrates a smaller sum of squares and a p-value (0.08311) above the conventional significance level of 0.05, implying that its influence on print time may not be statistically significant. Similarly, print speed exhibits a relatively low F value (1.7839) and a p-value (0.218415) above the threshold, indicating that it may not have a substantial effect on print time. These interpretations highlight the importance of considering layer height as a critical determinant of print time while suggesting that infill density and print speed may have a negligible impact in this specific context.

Consequently, it can be concluded that the input variables do not exert a significant influence on the dimensional variation of length. This conclusion is further supported by the regression analysis depicted in Figure 17-Figure 19, which consistently indicate a lack of significance. These results suggest that factors other than the selected input variables play a more critical role in determining the dimensional variation along the length or X-axis. Consequently, it can be inferred that the chosen input variables do not have a noticeable impact on the dimensions when compared to the original dimensions.

This indicates that none of the input variables has a significant impact on the dimensional variation of width. This lack of significance is further supported by the regression analysis depicted in Figure 21-Figure 22, which demonstrate a consistent pattern of insignificance. It becomes evident that variables217other than the selected input variables play a more significant role in determining the dimensional variation along the x-axis.The collected data indicates that, on average, the width deviates by approximately 0.5% from the originally intended width.

This finding is further supported by the regression analyses presented in Figure 23-Figure 25, which align with the results of the ANOVA tests.It is worth noting that the average height deviates by approximately 8.05% from the originally designed height. This notable discrepancy indicates an overestimation in height compared to the slight underestimations observed in the width and length dimensions.

The consistent overestimation in height is a significant observation, as it stands in contrast to the intended parameters established during the design phase.Table 8 ANOVA Analysis Table for Dimensional Change in Z DirectionEstimateStandard errorT valueP (>l t l)Infill density0.00155900.000855614.0680.106Layer height0.01111110.55889520.0200.985Print speed0.00096300.00372600.2580.803219Z n ie g n a h cl a n o i s n e m Di)m m(n o i t c e r i d3.45 3.4 3.35 3.3 3.25 3.2 3.15 3.1R² = 0.2916050 100 Infill density (%)150Figure 23 Infill Density vs Dimensional Change in Z Direction PlotZ n ie g n a h cl a n o i s n e m Di)m m(n o i t c e r i d3.45 3.4 3.35 3.3 3.25 3.2 3.15 3.10.1R² = 3E-050.15 Layer height (mm)0.2Figure 24 Layer Height vs Dimensional Change in Z Direction PlotZ n ie g n a h cl a n o i s n e m Di)m m(n o i t c e r i d3.45 3.4 3.35 3.3 3.25 3.2 3.15 3.1R² = 0.005945505560Print speed (mm/s)Figure 25 Print Speed vs Dimensional Change in Z Direction PlotAnalysis of Surface RoughnessTable 9 displays the examined input variables in the study, and it is evident that infill density holds the highest influence, explaining a substantial 78.06% of the observed variation, as depicted in Figure 26.

The significance of infill density surpasses that of the other two input220variables significantly. This finding is logical, as an increase in infill density results in a greater mass of iron material within the 3D print, ultimately leading to elevated surface roughness.In contrast, the analysis reveals that both layer height and print speed have minimal contributions, each explaining less than 1% of the overall variation in the data, as illustrated in Figure 27-Figure 28, respectively. Interestingly, it is observed that neither layer height nor print speed has a significant impact on surface roughness.

On the other hand, infill density emerges as the dominant and more influential factor, clearly indicating its strong association with surface roughness.Table 9 ANOVA Analysis Table for Surface Roughness MeasurementEstimateStandard errorT valueP (>l t l)Infill density0.0188840.9010335.4060.000641Layer height0.6400002.2817370.2800.786220Print speed0.0056220.0152120.3700.72127216.5)M µ (16s s e n h g u o r15.515e c a f r u s14.514R² = 0.780613.5020406080100120Infill density (%)Figure 26 Surface Roughness vs Infill Density Plot22116.5)M µ (16s s e n h g u o r15.515e c a f r u s14.514R² = 0.002113.50.10.120.140.160.180.2layer height (mm)Figure 27 Layer Height vs Surface Roughness Plot16.5)M µ (16s s e n h g u o r15.515e c a f r u s14.514R² = 0.003613.54550 55 Print speed (mm/s)60Figure 28 Print Speed vs Surface Roughness Analysis PlotConclusion and Future WorkThe analysis conducted in this study yields significant findings regarding the impact of various input variables on output responses in LCMMEX process.

Overall, it is evident that all input variables, namely infill density, layer height, and print speed, play a substantial role in influencing the output responses. Among them, infill density emerges as the most influential input variable, followed by layer height and print speed, respectively. Notably, these input variables have specific effects on different properties of the printed structures. In terms of magnetic properties, infill density significantly influences magnetic susceptibility, allowing for tailored magnetic properties in 3D-printed structures. Regarding thermal conductivity, layer height proves to be a critical factor, as smaller heights result in enhanced thermal conductivity by reducing air gaps and facilitating improved thermal transfer between layers. Surface roughness is significantly impacted by both infill density and layer height, with higher densities and smaller heights leading to smoother surfaces in AM.

Surface roughness is significantly impacted by both infill density and layer height, with higher densities and smaller heights leading to smoother surfaces in AM. Furthermore, infill density primarily determines the weight of printed parts, as higher densities correspond to increased weight due to greater material volume. Lastly, layer height plays a vital role in achieving dimensional accuracy, with smaller heights contributing to improved precision and consistency in the222dimensions of 3D-printed parts.

Lastly, layer height plays a vital role in achieving dimensional accuracy, with smaller heights contributing to improved precision and consistency in the222dimensions of 3D-printed parts. In conclusion, this study establishes the significance of all input variables in shaping the output responses across various aspects of LCMMEX process, providing valuable insights for optimizing the printing process and enhancing the quality of printed structures.For future work, exploring higher percentages of iron and sintering processes can be investigated to achieve higher material density, enabling applications in electromagnetic devices like actuators, transformers, and sensors.

veloping printable content: A repository for printable teaching modelsMary E. Knapp, Ryan Wolff, Hod Lipson Cornell University, Ithaca NY, USAAbstractAlongside the development of RP technology, there is an increasing need to develop and share printable content. Like digital photography and digital music, content drives technology as much as technology drives content. This paper describes the development and population of an open wiki-style website (3Dprintables.org) that houses an archive of printable models for education. Teaching models were chosen as the initial focus for this effort for two key reasons. First, quality educational models are difficult for teachers to obtain due to high prices, limited availability, and limited customization options; and second, many studies have demonstrated that learning is enhanced when students interact with physical models.

Such models are also indispensible tools for teaching the visually impaired and those with spatial reasoning difficulties. At present, the website contains models relevant to mechanical engineering, aerospace, biochemistry, mathematics, anatomy, and archaeology (e.g. proteins, airfoils, kinematics models, cuneiform tablets). These models are intended to serve as "seeds" to encourage educators to further develop and share printable models and the associated curricular materials.IntroductionWhile rapid prototyping/3D printing technology has advanced significantly in recent years, public interest has not kept pace. This is due in large part to a lack of printable content. Consumers who purchase or have access to a 3D printing system must design their own parts to print, which can be a time-intensive prospect and often requires experience with CAD software. Other emerging technologies in the past have faced similar challenges.

Other emerging technologies in the past have faced similar challenges. MP3 players did not become widely popular until large websites with many organized, downloadable music files were launched. Ease of content generation also contributes to the popularization of a technology. Digital photography content has been largely accelerated due to development of user-friendly image editing applications.This project is intended to address content generation for RP systems, with an initial focus on educational models. An open repository, 3Dprintables.org, has been constructed to contain an archive of free, testable, downloadable .STL files along with the source models and additional materials. Since the site is wiki-based, users can upload their own content to share freely as well as expand and comment on files already on the site.

Since the site is wiki-based, users can upload their own content to share freely as well as expand and comment on files already on the site. The site also includes tools that allow the novice user to immediately begin creating printable models from various sources such as data files, equations, and Protein Data Bank (PDB) files. The two main objectives of this website are increasing the amount of printable content freely available online and empowering educators to bring physical models back to the classroom by downloading existing models and designing their own.The widespread introduction of 3D printing to educational institutions has already begun. Prices for 3D printing systems are well within the reach of four-year colleges, community colleges,high schools and vocational schools and will continue to fall while the capabilities and speeds of these machines grow.

Indeed, many such institutions have already purchased 3D printing equipment and are integrating it into their existing curricula. The NSF recently funded the development of a rapid prototyping curriculum. RP technology is not only being used to supplement engineering and CAD courses but also as a means to introduce young students to engineering and increase enthusiasm for the field (Wohlers, 2005). This website will give educational institutions an additional use for the 3D printing systems they already own and perhaps encourage more schools to purchase this equipment.MotivationsPhysical models have always been an important teaching aid, especially in math and the sciences. Maria Terrell, a mathematics professor at Cornell, adamantly supports the development of models. “When handed a 2-dimensional representation of a 3-dimensional object many students have difficulty imagining what that object actually looks like, in many cases their interpretation is incorrect.

“When handed a 2-dimensional representation of a 3-dimensional object many students have difficulty imagining what that object actually looks like, in many cases their interpretation is incorrect. The image in the book is a level of abstraction of what the mathematical equation represents, models serve to bridge the gap and facilitate the learning of the student.” Traditionally, models were made by hand, often requiring a great deal of time and skill. Since these models were one-of-a-kind, only a small population of students could benefit from them at one time. Recently, increasing demands on teachers’ time have discouraged this type of model making. Older hand- crafted models are passed down within a school or department until they break or deteriorate. Also, these models are not able to evolve as scientific knowledge grows, making them increasingly inaccurate over time.Commercial models have emerged as an alternative to these hand-made models.

Also, these models are not able to evolve as scientific knowledge grows, making them increasingly inaccurate over time.Commercial models have emerged as an alternative to these hand-made models. Commercial models are generally more up-to-date than hand-made models and often more accurate. However, they are expensive, which can be prohibitive for schools with small budgets. For example, an anatomical model of the heart can cost up to $600. More complex anatomical models (a full torso, for example) can cost upwards of $3,000. The high cost of these physical models combined with the decreasing cost of simulation and visualization software and computer modeling have steered educators away from physical models. Teachers increasingly use computer models in place of physical models for classroom demonstrations. Additionally, commercial models do not yet exist for some classes, such as multivariable calculus. Teachers are forced to create their own demonstrations i.e.

Additionally, commercial models do not yet exist for some classes, such as multivariable calculus. Teachers are forced to create their own demonstrations i.e. a surface (the belly of a teddy bear) with tangent and normal vectors (pencils).While many students profit from computer models and simulations, blind students and students with spatial reasoning disabilities cannot benefit from these strategies. These populations are not insignificant. There are 10 million blind or visually impaired persons in the United States, and 93,600 of those are students. Out of these blind and visually impaired students, 55, 200 are legally blind (AFB, 2008). Students who are legally blind cannot benefit from illustrations in text books or on blackboards and can only interact minimally with computer simulations. Students who are visually impaired may not be able to experience the full value of these purely visual teaching strategies.

Students who are visually impaired may not be able to experience the full value of these purely visual teaching strategies. Additionally, students who are blind and deaf (10,800 in the US) are in even greater need of physical teaching models, since their dominant sensory experience is touch. Clearly there is a need for educational materials targeting these populations to ensure an equal opportunity for learning. Due to lack of availability, minimal customizability, and cost, manyeducators who work with visually impaired, blind, or deaf-blind students must construct their own models (Wentworth, 2008).Yet another group could benefit from the increased use of physical models – students with spatial reasoning deficiencies. These students do not perform as well as their peers in subjects such as anatomy which rely heavily on 2D illustrations of 3D objects (Rochford, K. 1985).

These students do not perform as well as their peers in subjects such as anatomy which rely heavily on 2D illustrations of 3D objects (Rochford, K. 1985). It has also been shown that male students outperform their female counterparts in spatial reasoning exercises (Stericker, 1982), perhaps because boys have more opportunity to practice those skills. Increasing the use of models in the classroom could help eliminate this disparity. Finally, teacher education benefits from the use of models. (Callison and Wright, 1993)Physical models have been shown to enhance learning in general student populations as well. Students learn in a variety of ways, and models allow students to include their sense of touch in the learning experience. “The role of experience is emphasized in Piaget’s description of cognitive development, that is, to know an object a subject must act on it and thus transform it - displace, connect, combine, take apart, and reassemble it.” (Cohen, 1983).

Science education especially benefits from the use of models.“If one goal of science education is to enhance and maximize an individual’s special conceptual ability, then access to manipulatives is advisable for those individuals. This access to manipulatives might also enhance development of their logical abilities… Internality is positively correlated with student achievement, and experience with manipulatives tends to move external subjects toward the internal end of the internal external continuum.” (Cohen, 1982)Graphic abilities have also been shown to have an effect on math and science performance.“A high degree of correlation was found between developed graphic abilities and high performance in mathematics, and a lower but still significant degree of correlation was found between developed graphic abilities and high performance in science.

The findings support the second hypothesis as well, suggesting that children with developed graphic abilities reported that their favorite subject was mathematics and their second favorite subject was science.” (Stavridou, 2008)Educational approaches that use models and manipulatives heavily (such as the Montessori Method) have demonstrated benefits for students. Montessori kindergarten students studied in Milwaukee, Wisconsin significantly outperformed their peers at traditional schools in standardized tests of reading and math (Lillard, 2006).The decreasing use of 3D models coupled with their broad usefulness in education was the motivation for creating 3Dprintables.org (Lipson, 2007). This archive site is intended to serve as a catalyst, encouraging educators to use models more widely in their classrooms and create their own models to share with other educators. Printed models are also cost-effective, especially for schools that already own a printer.

Printed models are also cost-effective, especially for schools that already own a printer. Material costs for most 3D printers range from less than $1 to $10/cubic inch for model material (less for support material). The printers themselves cover a large range in cost, from $3,000 to well above $40,000. Even the largest model on the website would likely cost under $50 (material only).Existing WorkRP technology is already being used to make educational models. A group at the University of Wisconsin–Madison is working with 3D protein models (Herman et. al., 2006). The group also runs a for-profit company that designs and sells the models. The Nanoscale Science and Engineering Center, also at the University of Wisconsin–Madison, has created 3D models to teach blind and visually impaired students about nanoscience (Greenberg, 2007). Some artists are also using RP technology to make their work more accessible.

Some artists are also using RP technology to make their work more accessible. Sculptor George Hart hosts a collection of .STL files for geometric sculpture on his website (Hart, 2008). Professor Gershon Elber of Technion hosts a collection of “impossible” Escher objects and scenes in 3D (Elber, 2008). For-profit sites also host large collections of models. The aim of this website is to collect and create educational models in one location to give educators easy access to a broad selection of free models.STL Generation ToolsMost commercial CAD programs will output a .STL file from a model (“How to Create Stereolithography (.STL) Files,” 2008). These .STL files can be sent directly to the printer to build parts or an assembly. Other commercial .STL generation tools exist for a number of applications, including making .STL files from medical imagery (“Medical Applications,” 2008), and freeform designs (“Freeform Modeling,” 2008).

We also created an in-house application capable of generating .STL files from data.Figure 1. STL-Generator screenshot. Screenshot of open-source utility for generating STL files from point, curve, and surface data, as well as molecular biology (PDB) files. This image shows renderings of a Lorenz attractor and a DNA segment. See (Lipson, 2008)STLGenerator is an open source utility for viewing and generating STL files from data. It can create an STL file directly from a list of points, line segments, surfaces, and PDB molecule files. The software uses the Marching Cube algorithm (Lorenson, 1987) to tessellate an iso-surface defined around a set of points or curves. A range of resolution can be selected resulting in finer or coarser meshes. The software is parallelized and makes efficient use of multiple cores. As an example, the external surface of protein molecule with 54,000 atoms can be rendered in approximately 5 minutes on a 2.8GHz quad core, resulting in a 1GB STL file.

As an example, the external surface of protein molecule with 54,000 atoms can be rendered in approximately 5 minutes on a 2.8GHz quad core, resulting in a 1GB STL file. A variety of options are available for generating partial models (selected chains in a large molecule), as well as adjusting atom size and curve widths, etc.The Website3Dprintables.org is a wiki-style site, hosted by WikiMedia. Any user with internet access can create an account, edit pages, and upload content. Wiki sites have been proven to work with non-technical users, which is consistent with the success and size of Wikipedia, the 7th most trafficked site on the World Wide Web (Desilets, A. 2005). Currently the models are divided into categories based on academic area, with some subgroups within these categories. Populated categories (so far) include kinematics, mathematics, molecular biology, archaeology, aeronautics, and chemistry.

Populated categories (so far) include kinematics, mathematics, molecular biology, archaeology, aeronautics, and chemistry. Also included on the site is a section on STL visualization and generation tools, including a free download of STLGenerator and instructions for its use. The page for each model contains images of the models and supporting images, information and suggestions on the educational use of the model, downloads for the .STL file and CAD, if applicable, links to resource material, and source and credit information.Creating the ModelsModels for this project were chosen in three ways. First, the authors chose models that are broadly applicable to high school or undergraduate students. These models relate to topics that are a staple of every high school science and math curriculum, such as revolutions about axes (calculus), catalytic enzymes and DNA (biology), and crystal structure of minerals (Earth science/chemistry).

Second, the authors searched the internet for models that had already been created in order to add these models to the collection for centralization purposes. Models obtained in this way include a Mayan pyramid (Figure 2-t) and the kinematic models from KMODDL (KMODDL, 2008, Figure 2 a-h)). Finally, the authors solicited input from educators (including Maria Terrell). Educators were asked to describe models that would be useful to them on a daily basis in their classrooms. This was an essential part of the project, since the key is for educators to take ownership in the site and begin creating their own models. It is simple to show an educator a few examples of the types of models that can be printed so he or she can see the limitations of the medium.

It is simple to show an educator a few examples of the types of models that can be printed so he or she can see the limitations of the medium. From there it is as simple as drawing a picture or giving the modeler an image from a textbook and modeling can begin.Mathematical models, such as the Lorenz Attractor (Figure 2-m), were generated from a list of XYZ points in a text file. In order to obtain the coordinate list, the appropriate equations were entered in MATLAB and a coordinate list was generated. This list was imported toSTLGenerator. The user can select whether STLGenerator should interpret the list as points, a curve, a surface or a solid volume. Then the user can choose the thickness of the curve to be generated. Revolved and intersecting solids were generated using a different method (Figure 2 j,k).

Then the user can choose the thickness of the curve to be generated. Revolved and intersecting solids were generated using a different method (Figure 2 j,k). A CAD program was used to create simple shapes and revolve them around axes.Models such as the revolutions about the axes and the intersecting volumes of cylinders can be used to teach the student how to visualize in addition to being the visual subject. By having a tangible model a teacher can move it around the axes, trace intersecting edges in three dimensions and otherwise show how the resulting volume was formed. This will help the student read the 2- dimensional representation (e.g. textbook illustration) more accurately in looking for key aspects. The more models a student sees the easier it is for them to visualize an unknown 3-dimensional object from a new equation or 2-dimensional representation.Selection of each protein to be modeled was based on several factors.

The first was the protein’s relevance to high school or introductory college level biology. Second, macromolecules with a reasonably small number of atoms (>60K) were selected for processing time considerations. This limit is not absolute, and should increase with faster processors and better printer software. Third, an effort was made to select models that had two parts – DNA and a binding factor for example – that could be printed separately and fit together, making the model interactive. The selected proteins were downloaded from the Protein Data Bank (PDB) (“RCSB Protein Data Bank”, 2008) and the PDB files were imported to STLGenerator. Separate “chains” within the molecule, such as a binding factor and the DNA it binds to, were rendered separately and exported as separate STL files. In PDB files, “chains” do not always refer to individual peptide chains. Proteins made up of several chains may appear in the PDB file as a single chain.

In PDB files, “chains” do not always refer to individual peptide chains. Proteins made up of several chains may appear in the PDB file as a single chain. Each chain was printed in a different color to highlight the interaction between them (Figure 2-q). A model T4 bacteriophage was generated in a CAD program. This model is not intended to be accurate to molecular detail, but rather demonstrate the general shape and mechanics of a virus. Ideally, the full virus could be rendered from a PDB file, but at present, virus capsid files contain <100K atoms and are very difficult to work with. Other possible large structures to be modeled include bacteria and eukaryotic cells or cell organelles.Aeronautical models were created in a CAD program. Several wings were produced with different purposes. A scale model of a Boeing 737 wing was made from four different airfoil profiles – one each for the root and tip of the wing and two mid-span.

Several wings were produced with different purposes. A scale model of a Boeing 737 wing was made from four different airfoil profiles – one each for the root and tip of the wing and two mid-span. The airfoil sections were obtained from the UIUC Airfoil Coordinates Database (Selig, 2008). This database contains thousands of airfoil sections, so the potential exists for many more wing models. The wing was tapered and scaled based on known Boeing 737 wing dimensions (“Wing Design”, 2008). Two other wings were generated for demonstration purposes. One uses a Clark Y airfoil, which is ubiquitous in general purpose aviation, and the other uses a generic high-lift airfoil. The purpose of these two wings is to demonstrate how airfoil shape affects wing capabilities. Both were tapered slightly and scaled to approximately the same size as the Boeing wing.

Molecular biology: (p) DNA strand, (q) Binding factor. Chemistry: (r) Buckyball, (s) Quartz (SiO2) crystal. Archeology: (t) Mayan pyramid, (u) Cuneiform tablet.Chemistry models were made in two ways. Simple geometric molecules (e.g. Buckyball, Figure 2-r) were generated in MATLAB and exported as XYZ coordinates. The XYZ coordinates were imported to STLGenerator and the atoms were scaled so that the representative spheres intersected enough for structural stability. Crystal structures were taken from text files (“Crystal Lattice Structures”, 2008). These text files were cleaned so that they only contained XYZ points. For crystals with two or more elements, radius scale factors were added in a fourth column to account for different atomic sizes (Figure 2-s). The modified text files were imported to STLGenerator and rendered. Some trial and error experimentation was needed to determine which atom scaling was best for structural stability and accuracy.

The modified text files were imported to STLGenerator and rendered. Some trial and error experimentation was needed to determine which atom scaling was best for structural stability and accuracy. A range between .6 and 2 was found to be optimal. At present, the software can only generate space fill models (intersecting spheres to represent atoms and bonds). This approach is useful because the resulting models are strong and stable and also best represent a real molecule, but some accuracy and clarity is lost, especially in larger crystals or molecules. Adjacent atoms that are not bonded may appear so because their spheres overlap. Ideally, space fill and ball and stick models should be used together for the most accurate picture of molecular or crystal structure. Very simple ball and stick models (H2O, sugars) may be created in CAD software.Another function of this site will be facilitating the exchange of archeological models between research institutions.

Very simple ball and stick models (H2O, sugars) may be created in CAD software.Another function of this site will be facilitating the exchange of archeological models between research institutions. Since many countries do not allow archeological artifacts to cross their borders, it can be difficult for archeologists to study artifacts from other countries, especially with shrinking travel budgets and increasing travel costs. There is also a great need to protect one- of-a-kind artifacts from damage, theft, or loss. Artifacts can be scanned with a 3D scanner and rendered in .STL format. These files can be posted on 3Dprintables.org for archeologists all over the world to download, print, and study (Figure 2-t). Increased printer resolution will also facilitate this effort.A person with no training in CAD software or 3D printing can easily create most of these models. STLGenerator allows a novice user to create complex molecular models in minutes.

STLGenerator allows a novice user to create complex molecular models in minutes. The key to this project is to empower educators to use these simple tools to customize models for their individual classroom needs. The website contains links to all the source material used for the models as well as all .STL files and CAD models. A list of STLGenerator settings and model dimensions accompanies each download. Users can easily recreate and customize any model on the site.FutureAs the costs of 3D printing systems continue to fall, colleges, high schools, and secondary schools will begin to purchase and use these systems. As the technology becomes more ubiquitous, educators will become increasingly aware of its power and begin integrating it into the classroom. This paper presents a jumping-off point for this projected trend.

This paper presents a jumping-off point for this projected trend. As awareness of this site and others like it grows, the art of model-making will be streamlined and placed back in the capable hands of educators.Acknowledgment This work was supported in part by the U.S. National Science Foundation CAREER award # CMMI 0547376.References Callison, P. L., Wright, E. L. (1993). “The Effect of Teaching Strategies Using Models onPreservice Elementary Teachers' Conceptions about Earth-Sun-Moon Relationships.” Proceedings of Annual Meeting of the National Association for Research in Science Teaching.Cohen, H.G. (1983). “A comparison of the affect of two types of student behavior withmanipulatives on the development of projective spatial structures.” Journal of Research in Science Teaching, 20(9), 875-883Cohen, H.G. (1982).

VELOPMENT OF 3D BIT-MAP-BASED CADAND ITS APPLICATION TO HYDRAULIC PUMP MODEL FABRICATIONTarou Takagi*, Tatsuro Yashiki*, Yasushi Nagumo*, Shouhei Numata* and Noriyuki Sadaoka*Power and Industrial Systems R&D Laboratory, Hitachi, Ltd.AbstractA novel 3D bit-map-based CAD has been developed. This program, named CellCAD,relies on techniques which provides easy manipulation of huge sized 3D bit-map, poly-triangle and 2D bit-map data. It can be attached to various kinds of additional plug-in processors toextend its functions and to customize it highly for broad applications. CellCAD can be appliedin fields using computed tomography digitizers and/or layering fabricators. The authors appliedCellCAD to hydraulic pump model fabrication. This paper reports the basic design concept andimplementation techniques of CellCAD from the viewpoint of design engineering, and alsointroduces the methods and results of an actual CellCAD application.1.

IntroductionRecently, many tools and methods for digital engineering are being developed and utilizedin manufacturing. To obtain digital shape data from real objects, high energy X-ray computedtomography (CT) is very useful, because it can digitize entire shapes including interior regionsaccurately and automatically in very short time [1].X-ray CT apparatuses obtain sectional images as two-dimensional (2D) bit-map data, whileconventional CAD systems handle boundary representation (B-rep) data [2]. B-rep data contain geometrical and topological definitions, called "features", which are not contained in bit-map data. Therefore, there is poor compatibility between bit-map data and B-rep data, and it is very difficult to convert bit-map data into B-rep data.The purpose of this study is to utilize piled-up X-ray CT sectional images, that is three- dimensional (3D) bit-map data, for various engineering activities.

The authors developed a novel CAD program, named CellCAD, which can hold and manipulate 3D bit-map data as engineering shape data. This paper reports the design concept and implementation techniques for CellCAD.4242. B-rep data and 3D bit-map dataThe difference in the two data formats, B-rep and bit-map, can be explained using Figure 1.A 3D B-rep datum is represented as a set of geometrical and topological definitions of shapeelements like vertices, edges, faces and bodies. On the other hand, a 3D bit-map datum isrepresented as a three-dimensional array of tiny cells known as voxels (volume picture cells).Each voxel contains inside/outside discriminant or relative density at its position.The greatest advantage of a 3D bit-map should be X-ray CT data acceptability. Sectionalimages obtained from X-ray CT apparatuses can be converted into 3D bit-map data by a piling-up method very easily.

Sectionalimages obtained from X-ray CT apparatuses can be converted into 3D bit-map data by a piling-up method very easily. Additionally, the 3D bit-map has other two significant advantages,"uniqueness" and "compliance". Uniqueness is the characteristic that the same shapes must berepresented with the same data. This is important for comparing two or more shapes to evaluatetheir sameness or to extract the difference between them. Compliance is the characteristic that the datum structure changes very little when a small deformation is applied to the shape. This isimportant for creating a shape by iterative computation like topology optimization.Even with these advantages, there are severe issues to deal with the large data size of 3Dbit-map. When representing even quite simple shapes, sometimes the 3D bit-map requiresmemory and storage capacity of 1 GB or more, while a B-rep requires less than 10 KB.

When representing even quite simple shapes, sometimes the 3D bit-map requiresmemory and storage capacity of 1 GB or more, while a B-rep requires less than 10 KB. Such ahuge amount of data is very difficult to hold, manipulate and display. This is the main reasonwhy all conventional CAD systems handle B-rep data.However, the situation has been changing recently.1. Computer performance has been drastically increased since the 1970s, when thefirst CAD system was developed. The latest PC models are able to handle hugeamounts of 3D bit-map data in reasonable times.V1:(1.0,1.0,1.0) V2:(1.0,0.0,1.0) V3:(0.5,0.0,1.0) V4:(0.5,0.5,1.0) E1:V1,V2 E2:V2,V3 E3:V3,V4 E4:V4,V5 F1:E1,E2,E3, E4,E5,E6 B :F1,F2,F3,F4, F5,F6,F7,F8E6V6V5E5E4E12F6F5V12E17V11V1F1E3F4E1V3E2F3V9E14V2E8V8Top layerBottom layerFigure 1. B-rep and bit-map formats425Voxels2. Many "layered manufacturing" tools have been developed.

B-rep and bit-map formats425Voxels2. Many "layered manufacturing" tools have been developed. Rapid fabricators and"voxel engineering" programs give 3D bit-map great utility in engineering fields.3.The quality and performance of X-ray CT has been greatly improved. 3D bit-map data containing even 2000·2000·2000 voxels can be obtained in a few days, from target objects with a 1000 mm diameter.Such significant changes have mitigated the weak points of the 3D bit-map, allowing itsadvantages to become the focus so there is a strong motivation for developing the 3D bit-mapCAD.3. Conceptual design of 3D bit-map CADAs the first step to develop a 3D bit-map CAD program, the authors started with theconceptual design. In order to determine its functional requirements, the authors made a list of issues to be considered. Comparison of 3D bit-map CAD with conventional B-rep CAD led toidentification of the following issues.1.Large data size.

Comparison of 3D bit-map CAD with conventional B-rep CAD led toidentification of the following issues.1.Large data size. A 3D bit-map datum containing 3000·3000·3000 voxels requires memory and storage capacity of more than 2.5 GB. This issue in turnleads to the following derivative issues.1-1. Network capacity. In the case of cooperative work by many operators, adistributed CAD system may cause network congestion due to largedata exchanges.1-2. Memory addressing limitation. Current Pentium-compatible 32-bitCPUs cannot address from a 4 GB memory space. Also the WindowsOS cannot address from a 2 GB memory space.1-3. Long process time. 3D bit-map CAD operators have to wait for a longtime until operations are completed. If the operators make a wrongoperation, they experience a terrific loss of time.2.Immature concept. The basic idea and methodology of engineering using 3D bit- map data have not been applied to many cases of practical production.

The basic idea and methodology of engineering using 3D bit- map data have not been applied to many cases of practical production. Therefore, requirements for 3D bit-map engineering tools have not yet been clearly defined. Functional changes or additions may be required in the 3D bit-map CAD program.3. Analytical usage. Conventional CAD operators know what they are handling, but this is different for 3D bit-map CAD operators handling X-ray CT data, because the target objects may not be their own products. The operators may want to use the 3D bit-map CAD for X-ray CT data analysis.426Many of these issues change the 3D bit-map CAD specifications. The authors found thefollowing solutions for the issues.1.PC cluster architecture (for issue 1). A PC cluster is a set of PCs connected by ahigh speed network. It is very useful for realizing scalable parallel processing.

A PC cluster is a set of PCs connected by ahigh speed network. It is very useful for realizing scalable parallel processing. Abit-map datum is a set of local information, therefore parallel processing is veryeffective to reduce process time.2. Client-server system (for issue 1-1). This consists of one server PC and one ormore client PCs. The server holds 3D bit-map data and processes them uponreceiving commands from the clients. In this system, large amounts of data arenot delivered to all clients to avoid network congestion.3.Partial loading and saving (for issue 1-2). Functions extracting partial data from3D bit-map files and functions inserting partial data into 3D bit-map files are4.required for handling large amounts of data over 2 GB. Pre-process and pre-load viewing (for issue 1-3). To avoid wrong operations, theresults are shown to the operators before they start the operations.

Pre-process and pre-load viewing (for issue 1-3). To avoid wrong operations, theresults are shown to the operators before they start the operations. Additionally, itis also better to show the results of threads, that is sequences including two ormore operations, before the operators start the threads.5. Operation describing language (for issue 1-3). To reduce the operating time,sequential execution of many operations is very effective. A special languageallows recording of operations and threads in literal scripts, modifying them usingtext editing programs, and executing them manually and automatically.6.Plugged-in functional modules (for issue 2). It is very effective to define theinterface between the kernel part and the plug-in modules to ease modification ofthe function set. Functions are attachable and removable when needed.7. Multi-viewer system (for issue 3).

Functions are attachable and removable when needed.7. Multi-viewer system (for issue 3). Many other CAD programs assign one viewerfor one datum, nevertheless, it is better to provide two or more viewers to easethree-dimensional shape recognition.8. Analytical functions (for issue 3). Extracting information from 3D bit-map data is very important to handle X-ray CT data from unknown target objects. Analysis, measurement and recognition functions are very useful for this purpose.4. Implementation of CellCADConsidering the results of the conceptual design, the authors have built a multi-purpose 3D bit-map CAD program named CellCAD.

Example CellCAD screen display1. Basic data management (data holding, memory management, display, etc.) 2. Basic user interface (command menu, information, help system, etc.) 3. 4. Plug-in module interface License control428The "CAD plug-in layer" consists of many CellCAD plug-in modules. They providevarious useful functions, some of which are listed in Figure 2.5. CellCAD applied to pump model fabricationCellCAD has been applied to hydraulic pump model fabrication. The target object was anexperimental model of a mixed flow pump impeller, designed with a 3D CAD program (Figure4). The diameter and the height were approximately 100 mm and 74 mm. The object wasfabricated by the selective laser sintering (SLS) method using metal material [3]. The SLSmethod, which is a rapid prototyping (RP) technique, included a furnace cycle to sinter thelaminated objects to give them high strength.

The SLSmethod, which is a rapid prototyping (RP) technique, included a furnace cycle to sinter thelaminated objects to give them high strength. The final material density was 7.9 g/cm3.This furnace cycle sometimes causes deformation of the objects, therefore a free-form shapevalidation method is required. For this purpose, the pump developer chose an X-ray CT apparatus (Hitachi HiXCT-6M). It worked very effectively to digitize the curved shape of the pump impeller into a set of 210 planes of sectional images, every one containing 900·900 cells. The total amount of the data was 162 MB.The shape validation was achieved with the following sequence. Views of the data duringthis sequence are shown in Figure 5.1.Input the sectional images. The sectional images were input and piled up to builda 3D bit-map datum (a).2.Input the designed shape. A poly-triangle datum, with the designed shape, wasinput from the 3D CAD program (b).3. Crop the 3D bit-map datum.

A poly-triangle datum, with the designed shape, wasinput from the 3D CAD program (b).3. Crop the 3D bit-map datum. The unused region around the object in the 3D bit-map datum was cropped to reduce thedatum size (c). The cropped datum contained 360·360·210 voxels.4. Adjust the scale of the poly-triangledatum. The 3D bit-map datum consisted of 0.35 mm-sized cubic voxels. The poly-triangle datum was resized to change its unit from mm to voxel.5. Obtain the geometric relationship between the poly-triangle and 3D bit-map data. The poly-triangle datum was overlaid on the 3D bit-map datum . The alignment could be done geometrically50mmFigure 4. Target pump impeller429acdand visually.6.Transform the 3D bit-map datum. The 3D bit-map datum was moved to thedesigned position, and rotated into the designed orientation (d).7. Obtain the density distribution.

The 3D bit-map datum was moved to thedesigned position, and rotated into the designed orientation (d).7. Obtain the density distribution. To know the relative density threshold whichseparates the internal and external regions, the relative density for the air part andthe metal part were measured, and their mean was taken as the threshold.8. Remesh the polygon facets. The facets, which are elements of the poly-triangledatum, were divided into smaller facets (e). This process is important todistribute the measured points closed enough together.9. Measure the surface offset. The original algorithm "virtual probe measurement(VPM)", which is a simulation of coordinate measuring machines, was applied toquantify the surface offset distribution between the 3D bit-map datum and thepoly-triangle datum. The accuracy of the VPM was higher than the voxel size. In many cases, it was within 0.1 voxels.

The accuracy of the VPM was higher than the voxel size. In many cases, it was within 0.1 voxels. Measured offset was converted into a colorand mapped on the facets.befg0.2 mm0.2 mmFailedFigure 5. Views of data during the shape validation sequence43010. Render the colored poly-triangle datum. The color-mapped poly-triangle datumwas rendered and the rendered images were pasted into a report which indicatedthe distribution of the deformation of the pump impeller (f, g).Significantly, all of these 3D bit-map operations are carried out within a day usingCellCAD. The report is very useful to analyze what happens during the furnace cycle and toimprove the furnace conditions.6. ConclusionsThe conclusions of this paper are as follows.1.

The report is very useful to analyze what happens during the furnace cycle and toimprove the furnace conditions.6. ConclusionsThe conclusions of this paper are as follows.1. A novel 3D bit-map CAD was developed, in order to utilize X-ray CT data and toenjoy the exclusive advantages of bit-map data.2.Issues for realizing the 3D bit-map CAD were listed, and ways to deal with them were found.3.The 3D bit-map CAD, CellCAD, was implemented as a 4-layer structure. It ispowered on client-server systems and/or PC cluster architecture.4. CellCAD was applied to hydraulic pump model fabrication. It is useful to analyzephenomena during the fabrication process of experimental models and to improvethe conditions for their fabrication.AcknowledgementsThe authors are very grateful to Prof. Noboru Kikuchi, the University of Michigan, for hiskind advice regarding this study. The authors are also very grateful to Mr.

Noboru Kikuchi, the University of Michigan, for hiskind advice regarding this study. The authors are also very grateful to Mr. Ichiro Harada, HitachiIndustries Co., Ltd., for giving them the chance to apply CellCAD to a practical problem.The authors are also very grateful to Mr. Ken Yamauchi and Mr. Kazuo Ohtomo, IbarakiHitachi Information Service Co., Ltd., for their unfailing cooperation.References[1][2]J. H Stanley, R. N. Yancey, CT-assisted Agile Manufacturing, Proceedings of SPIE, Nondestructive Evaluation for Process Control in Manufacturing (1996). J. D. Foley, A. Dam, S. K. Feiner, J. F. Hughes, Computer Graphics, Addison-Wesley (1990).[3] N. Sadaoka et al., Metal Selective Laser Sintering Technology Applied on HydraulicTurbine Design, The 8th International Conference on Rapid Prototyping (2000).431.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDevelopment of 3D Printable Part Library for Easy to Manufacture Components for Educational and Competitive RoboticsIndira Dwivedi, Rajeev Dwivedi, Bharat Dwivedi, Arun Rebbapragada, Arka RebbapragadaAbstractEducational and competitive robotics enable hands on learning and experimentation. Despite cost effective and ease of access of open source micro-controllers, drives and sensors, the structural components and brackets continue to be very expensive. Motivated by the Robotics for Everyone initiative, we are developing many easy-to-manufacture parts that will allow learners to easily 3D print parts for (1) Structural assembly of robot chassis (2) Sensor mounting (3) Electronic control mounting (4) Power supply (5) Various power drives.

The ecosystem of the robotic components is developed around extrusion structures and tubular elements and 3D printing is used for building the parts for testing and qualifying. Fixtures for mounting cameras for advanced machine learning and computer vision experiments are provided.IntroductionContextual understanding, aided by relevant examples and extensive hands-on experimentation, plays a crucial role in comprehending intricate concepts within subjects like science and math [1-4]. Educational Robotics, is one of the versatile platforms that connects classroom learning with real-life systems. It provides unique ability to verify and validate ideas through experimentation. Ardito et al. and Sana et al. [5,6] reported successful utilization of robots as educational tools in teaching intricate concepts in geometry, including ratios, estimation, and geometry. Their studies also highlight the practical relevance of incorporating robots in classroom learning.

Their studies also highlight the practical relevance of incorporating robots in classroom learning. Amico, Guastell, and Chella [7], claim that the integration of robots in learning experiences leads to improved comprehension of concepts and increased student engagement across different physics topics. Robots possess unique characteristics that set them apart from other machines, including multifunctionality, reconfigurability, and re-programmability [8]. For both learners and educators, reconfigurable equipment such as hobby and learning-grade robots have become a popular medium for learning and experimentation. Designing, constructing, and operating robots, therefore, require a diverse range of skills. Various aspects of robotics serve as instruments for conducting experiments in logical thinking, programming, and the principles of math and physics such as (but not limited to):1.

Various aspects of robotics serve as instruments for conducting experiments in logical thinking, programming, and the principles of math and physics such as (but not limited to):1. Robot navigation: Investigating the impact of acceleration, deceleration, velocity, anddisplacement relationships to achieve accurate and time-effective motion of a robot between two points in a 2D or 3D environment.2. Wheelbase correlation: Exploring the relationship between wheelbase, motor rotation, mobile3.4.robot wheel diameter, and the total distance covered while steering around obstacles. Lever systems and kinematics: Understanding the application of levers, gears, kinematic chains, and simple machines to control torque and force when lifting objects with a robot arm.

Lever systems and kinematics: Understanding the application of levers, gears, kinematic chains, and simple machines to control torque and force when lifting objects with a robot arm. Load distribution and center of gravity: Examining the principle of load distribution and the concept of the center of gravity when a robot traverses inclines and declines.2152The rising prominence of educational and competitive robotics has served as a catalyst in the growth of multiple commercial kits such as LEGO, VEX, REV etc. However, it is common for proprietary commercial robotics kits to cost several hundred dollars each. This poses a significant barrier for students from impoverished regions who lack the necessary resources to establish and finance a team.

This poses a significant barrier for students from impoverished regions who lack the necessary resources to establish and finance a team. Additionally, should a user wish to expand learning by expanding the scope of experiments and learning, related challenge is the dependency on specific manufacturers for computational platforms, actuation hardware (such as DC motors and servo motors), and structural components.Competitive robotics is another avenue where many educational institutions and community teams come together to participate and learn. Most of the competitions such as FIRST, VEX, WRO etc. are based on a prescribed kit provided by specific supplier. In addition to the robotic kit, the competitions require a robotics field that changes every season. Another challenge in competitive robotics is the need for exorbitantly priced field setup kits.

In addition to the robotic kit, the competitions require a robotics field that changes every season. Another challenge in competitive robotics is the need for exorbitantly priced field setup kits. Across various seasons and challenges such as FIRST, VEX, and similar events, the core task for robots typically revolves around picking and placing objects, often involving stacking or launching. Given the similarity of these fundamental tasks, it would be possible to reuse and reconfigure the same field elements, significantly reducing team costs. Alternatively, teams could employ a minimal superset of field elements that can be utilized for a range of challenges over multiple years. However, most leagues mandate the acquisition of new field elements annually, which can cost $1,000 USD or more per year [9-14].In contrast, many open source control and computation platforms like Arduino, Raspberry Pi, BBC Micro:Bit, BeagleBoard etc.

are highly popular in the DIY community and offer comparable capabilities. Moreover, there exists a wide array of affordable control, sensing, communication, and display modules that can be used in conjunction with these open-source platforms. These cost-effective kits have potential to become the go-to choice for numerous underprivileged schools and communities seeking to engage in STEM activities or participate in robotics competitions. The usability of these systems stems from their extensive flexibility and programmability. They provide a wide range of computational capabilities, multiple input/output ports, and seamless integration with numerous open-source libraries and electronic hardware. However, the learning with theses platforms will be limited to exploration in electronics and working principle by building circuits on a breadboard.

However, the learning with theses platforms will be limited to exploration in electronics and working principle by building circuits on a breadboard. As a Robotics learning platform, the electronic perception systems, the electronic actuators, programmable modules, navigation hardware, other functional platforms and end effectors are needed to build Entire systems. For example, principle of obstacle detection using ultrasonic sensing can be done by building a circuit and programming microcontroller; however, to enhance the learning the sensor needs to be mounted on a programable mobile robotic platform, where ultrasonic sensor detects an obstacle and provides input to a microcontroller that in turn drives a motor controller for a Robot to move, stop, slow down, or turn. Multiple robotic platforms based on open source programmable modules are available from any vendors [15,16]. However, the robotic platforms are fixed in configuration, limited to navigational and sensing platforms.

However, the robotic platforms are fixed in configuration, limited to navigational and sensing platforms. The platform cannot be expanded to test additional sensor and end effectors.Availability of parts and pieces, that allow easy reconfiguration to attach multiple sensors, actuators, end effectors, navigational platforms – differentially driven, actively steered, omnidirectional, legged, vision-based control is needed for comprehensive learning and experimentation. With such hardware, a learner may begin with basic odometry-based navigation experiments and gradually progress to more intricate tasks, such as machine vision and machine learning, all using the same hardware.

The hardware is also compatible with an effective mode of learning and progression [17,18].2153The research presented in this paper is inspired by Robotics for Everyone (R4E) initiative that enables accessibility of low cost, reconfigurable platforms for a comprehensive, experimentation and be competition ready within educational robotics. This work is a continuation of design and development work by authors to provide low cost experimentation platforms. They had earlier proposed flat pattern enabled brackets that can be manufactured using laser cutters or CNC routers [4]. Additionally, where possible, the brackets are designed as 2D flat patterns for ease of manufacturing. Many complex brackets are built by assembling a set of brackets manufactured from 2D flat patterns. In this research, we are presenting our findings in regard to using 3D printing as a means to build multiple cost effective educational Robotic platforms.

In this research, we are presenting our findings in regard to using 3D printing as a means to build multiple cost effective educational Robotic platforms. Each platform is built using a set of low-cost extrusions or tubular structures and brackets. The brackets can be arranged in different configurations to enable complex assembly of structural elements, functional elements as well as assembling sensing modules, power modules and multiple hobby grade microcontrollers. The repeatability of extrusion types, tubular members as well as brackets across various robotics kits ensures that a small number of parts suffice manufacturing of multiple robotic platforms and hence savings. The brackets can be 3D printed so that the learner does not have supplier limitations.In the following sections of the paper, different architectures of robotic platform along with the design of brackets is described.

We describe the features that are added to the 3D printable brackets for multiple use, ease of manufacturing and efficient material utilization. The paper concludes by outlining future work and proposing a mechanism for learners to access and share the available resources including 3D printable parts and programming library and codes.Path of learning and advancement in educational RoboticsFigure 1 : Path of progression for learning with Robotic PlatformA common pathway for learning with educational robotics involves initially experimenting with physical systems and then identifying opportunities to introduce complexity, autonomy, and the need for programmability. This approach naturally establishes a context for logical thinking and programming. The design of the Robotic Platforms is driven by the same intent.

This approach naturally establishes a context for logical thinking and programming. The design of the Robotic Platforms is driven by the same intent. As described in the Figure 1, the kit will allow building of a simple robot that can be controlled differentially for a feedforward odometry based path planning, then environmental perception using sensing is added to the robotic platform. End effectors are added to the robot to performing different tasks. Later, features to add vision systems and machine learning modules are added to the robot.As described in Figure 2 , Figure 3 and Figure 4 we start with simple platforms such as differentially driven Robotic Platform and add additional modules as the learner ventures further into autonomy by adding Robot navigation. Additionally, the complexity of the robot is enhanced by adding additional end2154effectors, complex drives such as omnidirectional mobile modules and industrial architecture.

Additionally, the complexity of the robot is enhanced by adding additional end2154effectors, complex drives such as omnidirectional mobile modules and industrial architecture. The robots are built using 10mmX10mm extrusion and Metric 3mm (M3) hardware is used for the system assembly. The brackets for assembly of the extrusions, motors, sensors, microcontroller installation are 3D printed parts.Figure 2 : Mobile Robots and increasing complexity (CAD Model)Figure 3 : Additional Industrial Robot Architectures built using 3D printed brackets and 10mmX10mm extrusionsFigure 4 : Mobile Robots built using 3D printed brackets and 10mmX10mm extrusionsDesign for Accessibility2155During the development of the kit, we adhered to several key principles. These include:Cost Consideration: The total cost including the structural elements and the brackets should not exceed 75 USD. This cost does not include the controller and sensors.

These include:Cost Consideration: The total cost including the structural elements and the brackets should not exceed 75 USD. This cost does not include the controller and sensors. To enable autonomous navigation of the robot following components should be added: ultrasonic sensors, IR reflective sensors, a controller, and a battery kit. The suggested cost for additional modules does not exceed 30 USD. Therefore, the total cost for an autonomous microcontroller-based Robot will not exceed 105USD. For advanced learning and experimenting in vision and machine learning a camera can be added and the microcontroller can be replaced with microcomputers such as Raspberry Pi. The total cost of a vision enables Robot (Including mobile robot platform, motors, camera and microcomputer) will not exceed 125USD.Manufacturability: The parts can be produced using most personal desktop 3D printers and easily accessible machines.

Both the structural features and assembly hardware features can be printed without requiring additional processing.Interchangeability: The parts are designed to be interchangeable and compatible with a wide range of open-source controllers and modules.Alternate designs using Flat Stock Material: Should an organization or individual wish to mass produce the parts, the 3D models can be manufactured by stacking a set of flat structures. The flat structures can be manufactured in a very cost-effective manner using laser cutter or CNC-mill. In cases where flat patterns are limiting, a combination of flat patterns is used to simulate 3D features. The authors wish to refer to their previous work on the same [4].Reusability and Reconfigurability: As described in Figure 2 and Figure 3 the core framework for the mechanical hardware is built using 10mm x 10mm T-slot aluminum extrusions or tubular elements.

To further enhance versatility, lightweight extrusions can be easily customized for different structural and functional purposes using a range of brackets.Figure 5 : CAD models for some brackets as designed for Extrusion vs Tubular frameworkWe extensively utilized computer-aided design (CAD) software to design and test a variety of attachments during the initial concept phase. Figure 5 showcases a selection of these designs, which can be 3D printed and assembled with the structural elements. A learner can download and 3D print these2156models from the Robotics for Everyone website [19]. We used Solidworks (Rev 2021) to design the parts. The CAD files are available in original Solidworks source format (SLDPRT for parts and SLDASM for assemblies). Where applicable, the flat patterns are provided in DXF format. The STL files are provided so that user can 3D print the files directly.

Where applicable, the flat patterns are provided in DXF format. The STL files are provided so that user can 3D print the files directly. For the CAD users other than Solidworks, STEP files are provided.These extrusion framing elements feature continuous T-slots that allow for easy attachment of various types of fasteners. The rails have slots on all sides, providing maximum design flexibility. Unlike many other structural elements that require pre-drilled holes for assembly, T-slot aluminum framing allows for the insertion of T-nuts or bolts to create assembly features. This feature eliminates the need for fixed hole positions and enables assembly in almost any increment.Figure 6 : Cost effective Robot built using tubular elementsAnother readily available structural element is tubular structures. For an extensive set of experiments in educational robotics and compatibly with DIY (Do It Yourself) electronic parts we are providing robotic platforms for 1/8 tubing (Figure 6).

For an extensive set of experiments in educational robotics and compatibly with DIY (Do It Yourself) electronic parts we are providing robotic platforms for 1/8 tubing (Figure 6). The tubing is easily available for PVC, aluminum and steel. The tubing can be purchased for low cost in segments of 5ft incremental lengths. The tubing can be cut as needed. One advantage of the tubing is that many off the shelf fittings are easily available. Additionally, a library of 3D printable parts for structural-, sensor-, drive- brackets are provided (Figure 5). In addition to the structural assembly elements, we have also designed brackets specifically for mounting functional elements such as motors and sensors.

In addition to the structural assembly elements, we have also designed brackets specifically for mounting functional elements such as motors and sensors. In addition to the brackets for extrusion elements, the brackets for tubular element are available for download, allowing a learner to easily incorporate additional components into robotics projects.Considerations for 3D printable elementsThe experimental robots must withstand static and dynamic loading. A learner may reconfigure the robot for multiple experiments, therefore the versatility of bracket design will reduce the part count and keep the overall manufacturing cost low. Additionally, each bracket should be designed for multiple cycles of assembly and disassembly.The Metric 3mm (M3) hardware is used for the system assembly. Two types of holes are designed for system assembly (1) threaded (2) clearance. The holes that require threads for assembly, are2157manufactured for without thread with interference fit diameter.

Two types of holes are designed for system assembly (1) threaded (2) clearance. The holes that require threads for assembly, are2157manufactured for without thread with interference fit diameter. When a screw is assembled into the holes, the screw taps and creates a thread. Sensors that are based on reflection from the robot environment (e.g. floor), require accurate mounting. Additionally, vibration isolation is required for inertial and reflective sensors. To reduce the measurement errors, robot structure shall provide a stable datum to mount the sensors.Table 1: Failure test for reassembly of partsFigure 7 : Two types for damages seen due to multiple cycles of assemblyWe manufactured the parts using different nozzles used within hobby grade Open-source desktop 3D printers. The manufacturing was done one Ender 3 pro (Fused Filament Fabrication) FFF machine with PLA+ filament. We performed basic test for fatigue and structural integrity.

The manufacturing was done one Ender 3 pro (Fused Filament Fabrication) FFF machine with PLA+ filament. We performed basic test for fatigue and structural integrity. For the brackets that at are cantilevered, we add structural features such as ribs. A detailed analysis is not within the scope of the research presented in this paper. However, we are reporting some of the testing and findings based on basic manual assembly. We tested the designs for 10 reassembly cycle for structural thickness of 3mm. Each test was performed for approximately 5 similar parts. The speed of extrusion is adjusted for the nozzle size. For the parts manufactured with nozzle size 0.8mm or above, the assembly feature such as screw holes start to crack after 5 cycles of reassembly. The smaller nozzle diameters also take longer time to build. As a rule of thumb, we recommend using 0.4mm nozzle diameter for 3D printing. The learner may choose their own parameter based on the choice of their printer and material.

As a rule of thumb, we recommend using 0.4mm nozzle diameter for 3D printing. The learner may choose their own parameter based on the choice of their printer and material. Should significant reuse and reconfiguration of the 3D printed parts required learners may perform a quick experiment to determine the parameters amenable. Table 1 summarizes results of our experiments. Figure 7 describes two types of damages seen during the test. In some cases, the crack will lead to complete dislodging of a segment. Whereas in some other cases we saw crack develop. A partly cracked bracket may continue to function as needed as long as not a significant static or dynamic load is not applied. However, we strongly recommend that sensor brackets should be immediately substituted,2158since the position of sensor with respect to environmental features is critical for the robot to function.

Additionally, we found that irrespective of the nozzle diameter, the threads introduced by assembling hardware in interference fit holes do not show any deterioration for thicknesses tested.Our estimates of the price of filament material consumed for different brackets range from 0.15USD (United States Dollar) to 0.85 USD. The price estimate of the linear extrusion is approximately 0.11USD and 0.05USD for tubular structures. The cumulative price estimate for a complete navigational platform with electronic controls, drives and motors is approximately 45USD.The part library is done for in-school learning and experimentation. The part library will enable complete learning and experimentation per path of progression described in Figure. For learners who wish to compete, the part library will allow building a robot that will meet FIRST, VEX, WRL, BEST and WRO form fit specifications for a fraction of cost (115 USD vs approximately 500USD).

However, the robot will not be admissible for FIRST and VEX leagues due to league prescribed elements such as LEGO and VEX.Test Cases for 3D printable partsFigure 5 describes a representative set of parts from our CAD library as applicable for extrusion-based platform and tubular platform. The models are available in STL format to 3D print it as is. Additionally, the models are provided in STEP format should the user want to customize the models per specific needs. We are describing below a sample of parts that enable reconfigurability and multifunctionality.Part manufacturing for reconfigurability and multifunctionality:Figure 8 : Bracket for assembly in multiple configurations and reuseFigure 8 describes a bracket used for connecting extrusion structure at 90 degrees. Additionally, the bracket has features being reconfigurable and multi-functionality.

Additionally, the bracket has features being reconfigurable and multi-functionality. The bracket has a feature that will allow assembling a third extrusion to build a 3D frame.2159Figure 9 : A plate to enable installation of multiple Open source microcontrollers, computers, drivers and power supplySimilarly, as described in Figure 9 the mounting plate for the microcontrollers and battery is designed with assembly features to enable assembly of multiple open source microcontrollers (Ardunio Uno, Arduino Mega) and microcomputer such as Raspberry Pi on a common plate.Figure 10 : Sample 3D printed parts used for building RobotsFigure 10 describes brackets for installing popular sensors for environment perception. The bracket allows installing the ultrasonic sensor in front as well as rear side of the robot. Additionally, using the same bracket sensor can be installed along the side to enable wall following.

The bracket allows installing the ultrasonic sensor in front as well as rear side of the robot. Additionally, using the same bracket sensor can be installed along the side to enable wall following. Reflective sensors are used for line following and similar environmental perception. The sensitivity of the sensor may vary due to manufacturing variation as well as the robot environment, therefore it is critical to have some2160adjustability to the distance of the sensing end effector to the surface. The bracket as described in Figure 10, enables installation of the sensor with the adjustability for height.The limitation of material availability manifests in mixing two different type of structural platform forFigure 11 : Adapter for connecting the extrusion-based platform to other commercially available Robotics Platformbuilding robots. The learners may expand the capability of robot by interchanging components between extrusion and commercial platforms that use channels.

The learners may expand the capability of robot by interchanging components between extrusion and commercial platforms that use channels. In certain cases, a learner might choose to design and produce a robot using one kit, but opt to incorporate other subsystems, such as drives, from a different system. This decision could be driven by either the advantage of the form factor or the availability of specific components. Adaptive brackets may be used (Figure 11). A user may design the adapting bracket per need. The adapter, as described is 3D printed to assemble parts from popular channel-based system in two different configurations. The part is manufactured by Fused Filament Fabrication, using Orange Acrylic Styrene Acrylonitrile (ASA) printed with T40 tip.Field Elements for Robot environment.Figure 12 : Sample field elements built using cardboard and tape -1.

(The plat pattern is provided)Robotic experiments require interaction of the robot with different types of environments for navigation and process planning. The robot may be programmed to move from one point to another and solve2161different tasks in feedforward mode or interact with the environment using sensors to take decisions. The robot environment elements may be simple obstacles or complex three-dimensional elements. Alternatively, the robot elements may be systems that can be actuated by engaging with the Robot. To keep the cost for building field elements low, we are providing schematics and details so that easily available material such as tape, paper, wire and carboard can be used to build field elements. As described in Figure 12, CAD tools provide utilities to create flat-pattern to build the parts out of cardboard. However, accuracy of the shape as well as accurate placement of field elements is required.

However, accuracy of the shape as well as accurate placement of field elements is required. 3D printed parts can be used for accurate placement and assembly. Additionally, as described in Figure 12, 3D printed parts can be combined with cardboard based part to build and place field elements accurately.Models and coding libraryA survey encompassing 147 recently published studies, [20] suggests that programming the robot provides additional benefits including (1) enhancing their understanding of abstract concepts, (2) providing a feedback-oriented learning platform, (3) fostering collaboration with peers, and (4) offering opportunities for exploring and gaining deeper insights into real-world problems. To allow ease of entry, many basic programs are made available by Robotics 4 Everyone as a library. The coding library as provided is created in C++ for Arduino platform. The details of programming libraries and functions are beyond the scope of this paper.

The coding library as provided is created in C++ for Arduino platform. The details of programming libraries and functions are beyond the scope of this paper. However, we would like to introduce the basic coding model of how it is made available to the learners to integrate and experiment.Figure 13 : Pictorial representation of the Coding libraries and hierarchyAs described in Figure 13, basic DC motor control is using Pulse Width Modulation (PWM) and is made available as basic function. The Robot navigation library is implemented as a combination of motor controls including the speed and direction. The sensing functions can be combined with navigation functions for higher level robot autonomy. This approach allows learner to dive deep and explore the programmability of the robot. Pseudocode described in Appendix provides the details of navigation implementation.

This approach allows learner to dive deep and explore the programmability of the robot. Pseudocode described in Appendix provides the details of navigation implementation. The modules may be combined with input from sensor and potentially vision-based machine learning to control a Robot.Conclusion and future workThis paper outlines the utilization of 3D printing to manufacture brackets and elements that can be coupled with cost effective extrusion and tubular elements to assemble different types of robotic platforms. Additionally, 3D printing can be used to manufacture field elements for Robot environment.2162The models are provided online for the learner community to download and use to build cost effective robots. Design of parts is done for reusability of parts. The designed elements are used for structural, functional and accurate mounting of sensors.

Design of parts is done for reusability of parts. The designed elements are used for structural, functional and accurate mounting of sensors. Additionally, source files are provided for users to modify when need basis.As a pilot program, 5 sets of kits were made available to two middle schools in one of the financial back ward regions of North Eastern India. Each school was provided with (1) Arduino based kit along (2) extrusions set and (3) a 3D printer each. Across 3 months of learning and experimentation many students were able to implement differential robots along with passive as well as active end effectors. The pilot concluded with a competitive event were teams in group of 2-3 student each implemented pick-up and place challenges. The educators suggested that the cost effectiveness was helpful in their ability to offer the program to multiple students. The comparative price of off-the-shelf robot kits such as Lego spike prime, VEX were not affordable.

The comparative price of off-the-shelf robot kits such as Lego spike prime, VEX were not affordable. Authors believe that free availability of resources and models will enable additional communities to benefit from this initiative.Ability to build and test various functional units such as different types of gears and kinematic-linkage was limited. The functional elements such as gears when produced with hobby grade 3D printers lack tolerance and do not function smoothly.

velopment of a 3D Printing Method for Production of Dental Application Li Yanga, Shanshan Zhanga, Gustavo Oliveirab, Brent Stuckera aDepartment of Industrial Engineering, J.B. Speed School of Engineering bDepartment of General Dentistry and Oral Medicine, School of Dentistry University of Louisville, KY, 40292Abstract Traditionally, the manufacturing of dental restorations, including crowns, veneers and other structures made by ceramics, is a labor‐intensive and time consuming process. Additive manufacturing has the potential to significantly decrease the time and cost associated with this process. This work performed preliminary investigation for the feasibility of dental restoration parts printing using the ExOne M‐Lab system with a commercialized dental porcelain powder. The porcelain powders were characterized, and two measurements, including pre‐sintering and addition of flow agent, were taken in the attempt to improve the processability of the original powder feedstock.

The results showed that while the addition of flow agent has more significant effects in improving the flowability of the powder used, the post sintered parts exhibit considerable shrinkage and residual porosity that necessitates further investigation.Introduction Over the past decades, there has been a trend of using metal‐free restorations in the dental field (Denry and Holloway, 2010, Guess et al., 2011). Ceramic dental restorations have been increasingly used due to their outstanding aesthetic features and resistivity to chemicals (Conrad et al., 2007, Kelly, 1999). Traditionally, the dental ceramics were produced by hot‐pressing, sintering or slip casting processes (Denry and Holloway, 2010), which in general lack sufficient flexibility in the accurate customization of the parts. Machining was used as well, but often suffers from the high hardness and low toughness of the ceramic feedstock.

Machining was used as well, but often suffers from the high hardness and low toughness of the ceramic feedstock. In recent years, a CAD/CAM based method was developed to produce dental restorations with geometries accurately matching that of the patients’ (Filser et al., 2003, Guess et al., 2011). The basic procedures of this process involves the pre‐sintering of ceramic blocks followed by a CNC machining process that produces accurate shapes, then the green machined parts are subject to a densification sintering process to generate the final densities as well as mechanical performance. The use of the two‐stage sintering helps reduce the issues associated with machining of hard ceramics, therefore increasing the efficiency of the process with lower cost. However, due to the shrinkage in the second sintering stage, the accuracy of the machining could be significantly reduced, which compromises the benefit intended to be provided via machining.

However, due to the shrinkage in the second sintering stage, the accuracy of the machining could be significantly reduced, which compromises the benefit intended to be provided via machining. In addition, the relatively complicated procedures also result in longer manufacturing time and higher cost.With the ability of manufacturing parts directly from a CAD model with adequate accuracy and minimal waste, additive manufacturing holds great potential for the future production of custom dental restoration parts. Several laser sintering based studies have been presented in the attempt to fabricate dental restorations with various materials (Li et al., 2000, Hagedorn et al., 2011). The use of a laser for the directsintering of ceramics enables one‐step processing for the manufacturing of these parts, and provides the potential for direct manufacturing of ceramic‐over‐metal dental parts.

On the flip side, thermal cracking is a common issue for these processes (Wilkes, 2010), and the feature resolution is also often limited due to the partial sintering of the surrounding ceramic powders (Hagedorn et al., 2011). Recently, a direct write based printing process was investigated for the fabrication of zirconia dental prostheses (Ebert et al., 2009). The process selectively deposits a zirconia based suspension on graphite substrates, followed by a post sintering process to achieve the final density. The resulting parts exhibit high geometrical accuracy and density, as well as good mechanical strength. In this work it was also reported the potential issue with the nozzle clogging (Ebert et al., 2009).In this study, the ExOne M‐Lab was used in an attempt to fabricate ceramic parts from off‐the‐shelf commercial ceramic powders used for dental applications.

While the system is relatively new, there have been studies that utilize binder jetting to make ceramic parts (Cima et al., 1995, Uhland et al., 1999). This process offers some potential advantages in ceramic printing, such as the flexibility with different ceramic materials, the relatively high feature resolution, and easy process control; therefore binder jetting was adopted for this study with future developments in mind. The goal of this study is to identify the feasibility of the system for quick and accurate fabrication of complex‐ shaped, 3D dental restoration parts. Based on the preliminary study, future work can be determined in the effort to produce high quality parts with minimal defects.Materials and Methods The primary material used for this study is VITA VM13 Base (Vita Zahnfabrik, Germany), which is a leucite‐reinforced glass ceramic widely used for dental practice.

The chemical composition and basic properties of the powder provided by the manufacturer is shown in Table 1 (Vita VM‐13, 2009).PowderChemical composition (wt%)Particle sizeLinear coefficient of thermal expansionFlexural strengthVM‐13 BaseSiO2: 59‐63%, Al2O3: 13‐16%, K2O: 9‐11%, Na2O: 4‐6%~18µm13.6‐14x10‐6 K‐1~120MPaTable 1 General information of VM‐13 BaseThe original powders exhibit significant aggregation which made it unsuitable for the process. Preliminary particle size analysis (Microtrac S3000) indicated that the characteristic size of the aggregation was around 100‐105µm, as shown in Fig.1. Two factors were considered to contribute to the aggregation. From Fig.2, it could be seen that the powders have a very irregular morphology and a rather wide size distribution, which could significantly reduce its flowability. In addition, the large surface activity ofthe silica powders also makes them prone to clumping (Bagwe et al., 2006).

In addition, the large surface activity ofthe silica powders also makes them prone to clumping (Bagwe et al., 2006). Therefore, Pre‐sintering and flow agent addition were implemented in an attempt to improve the issue.Fig.1 Particle size analysis results for the VM‐13 Base powderFig.2 Morphology of the original VM‐13 powderThrough a preliminary trial it was found that at about 600°C of sintering temperature, significant densification does not occur for the powder. Therefore, 600°C was chosen to be the pre‐sintering temperature in order to achieve fast effects. The holding time for pre‐sintering was kept at 30 minutes, followed by furnace cooling. Surface‐modified R972 SiO2 powder (COSMOS Plastic & Chemicals) was used as the flow agent. The powder is composed of >99.8% fumed silica treated with dimethyldichlorosilane (DDS), with an average particle size of 16nm.

The powder is composed of >99.8% fumed silica treated with dimethyldichlorosilane (DDS), with an average particle size of 16nm. The flow agent was added to the original VM‐13 powder and mixed by hand until satisfactory improvement of flowability was observed. The flowability was evaluated by the angle of repose for the original and treated powders. In addition, micrographs were taken for the pre‐sintered powder to evaluate the evolution of powders. After the treatment, the processable powders were used toprint sample crown parts using the M‐Lab. The binder used for the process was the ExOne PM‐B‐SR1‐04, an ether solvent based binder. After printing, the green parts were dried in the oven at 70°C overnight. The schematic of the post sintering process is shown in Fig.3. The green parts were held at 500°C for 1 hour to burn out the binders, followed by a two‐step sintering sequence at 700°C and 850°C, respectively.

The green parts were held at 500°C for 1 hour to burn out the binders, followed by a two‐step sintering sequence at 700°C and 850°C, respectively. The dimensions of the final parts were measured with caliper before and after sintering, and results were compared to evaluate the shrinkage. The cross sectional microstructure of the sintered samples were also prepared and observed using optical microscopy.Results and discussion Pre‐sintering did not seem to have significant effect on the flowability of the powder. Fig.3 shows the morphology of the powders after the pre‐sintering. From Fig.3 it could be clearly observed that significant sintering occurred which resulted in the generation of large powder aggregations. On the other hand, the flow agent had a more significant effect on the flowability of the powder at the ratio of approximately 7% in volume. Fig. 4 shows the angles of repose for each sample, and Table 2 shows the results of the measurement.

Fig. 4 shows the angles of repose for each sample, and Table 2 shows the results of the measurement. Pre‐sintering was used for ceramic powders with fine particle sizes in previous studies using selective laser sintering (Harlan et al., 1999). However, from this study it seems apparent that with highly irregular powders, even if the pre‐sintering could reduce the total surface area, the increase of particle irregularity will cancel out all the benefits and potentially result in the reduction of flowability.

It was expected that the addition of flow agent could improve the overall powder flowability by serving as a lubrication interface between the large VM‐13 powders, which was verified by the study.Fig.3 Morphology of VM‐13 powder after pre‐sintering(a) Original(b) Pre‐sintered Fig.4 Angles of repose for each type of powder(c) Flow agent additionPowder Original Pre‐sintering 7%v addition of flow agentAngle of repose (degree) 64 66 56Table 2 Angles of repose for each type of powderThe powder with flow agent addition was successfully used for the printing of crown parts in the M‐Lab and sintered subsequently. Fig.5 shows the original part as well as the post sintered parts. Significant shrinkage could be observed. The dimensions as shown in Fig.6 were measured for three samples, and the results are shown in Table 3. The results are quite consistent for each dimension, and an average of 25%‐32% shrinkage was observed in each dimension compared to the green parts.

The results are quite consistent for each dimension, and an average of 25%‐32% shrinkage was observed in each dimension compared to the green parts. The shrinkage values are considerable but reasonable considering that the powder is highly irregular. A slight difference of shrinkage between the build direction (aligned with the dimension H) and the other two directions were observed, which could be caused by the burnout of the binder. There is also a slight difference of shrinkage between the two planar directions perpendicular to the build direction, as could be seen from Table 3. This could be associated with the specific geometry of these samples such as wall thickness. However, without further investigation, it is difficult to determine the likely cause of this phenomenonBeforeAfterFig.

However, without further investigation, it is difficult to determine the likely cause of this phenomenonBeforeAfterFig. 5 Green and post sintered crown partsFig.6 Measurement of dimensionsSampleL (mm)W (mm)H (mm)Before sintering1 2 311.50 11.45 11.4311.55 11.55 11.586.00 6.07 6.03Post sintering1 2 38.80 8.61 8.608.16 8.11 8.094.12 4.08 4.13Linear shrinkage1 2 3 Average23.48% 24.80% 24.76% 24.35% Table 3 Shrinkage of the samples29.35% 29.78% 30.14% 29.76%31.33% 32.78% 31.51% 31.88%Fig.7 shows the microstructure of the ceramic samples after sintering. In general the sintering seemed to have achieved a relatively homogeneous densification. Large voids as well as some light particles could be clearly observed. The size of the voids varied butwere at the magnitude of 50‐100µm. This seemed to have a correlation with the size of the aggregation of the original powder.

The size of the voids varied butwere at the magnitude of 50‐100µm. This seemed to have a correlation with the size of the aggregation of the original powder. One possible cause could be that during the powder spreading process, the large aggregation was displaced by the roller due to the temporary adhesion to the roller surface, which resulted in surface voids on the newly spread layer. Another potential source that could contribute to the void formation is the addition of the flow agent. The mechanism of the surface modification in the flow agent was to create a hydrophobic surface on the silica particles. Since sufficient wetting is required to form a continuous bonding between the binder and the powder, the existence of the flow agent could potentially affect the continuity of the green part, leaving voids that become defects after the subsequent processes. Further studies are needed to identify the potential impact of the flow agent in the final densities of the parts.

Further studies are needed to identify the potential impact of the flow agent in the final densities of the parts. In addition, the use of 850°C as the densification sintering temperature was largely due to the limitation of the furnace, while in the reference instruction for the VM‐13 powder (Vita VM‐13, 2009), it was recommend that the green part be sintered at 920°C. The lower temperature could also contributed to the porosities observed in the sample.Fig.7 Microstructure of the post sintered partConclusion In this study, an off‐the‐shelf dental veneer powder was successfully printed using an M‐ Lab. The original powder had significant aggregation issues, and two measurements were used in the attempt to improve the flowability of the powder. Addition of the flow agent of about 7% volume was shown to be an effective method to improve the original powder flowability.

Addition of the flow agent of about 7% volume was shown to be an effective method to improve the original powder flowability. At relatively low sintering temperature, the final parts showed homogeneous densification and continuous microstructures, although large voids werepresent in the final parts. The shrinkage of the parts after sintering was around 25‐30% in each directions, and anisotropic shrinkage was observed in all three directions. Further investigations are needed to identify the cause of the anisotropic shrinkage as well as the source of the large voids in the final parts, and additional mechanical testing is also required to further characterize the quality of the process.Reference R. P. Bagwe, L. R. Hillard, W. Tan. Surface Modification of Silica Nanoparticles to Reduce Aggregation and Nonspecific binding. Langmuir. 22(2006): 4357‐4362. M. J. Cima, J. Yoo, S. Khanuja, M. Rynerson, D. Nammour, B. Giritlioglu, J. Grau, E. M. Sachs.

velopment of a Binder Formulation for Fused Deposition of CeramicsThomas McNulty, Ivan Cornejo, Farhad Mohammadi, Stephen C. Danforth, and Ahmad SafariRutgers University Department of Ceramic and Materials Engineering Piscataway, NJ 08854AbstractA new binder formulation has been developed for Fused Deposition of Ceramics (FDC) which consists of commercially-available polymer constituents.. This formulation was used. in conjunction with lead zirconate titanate (PZT) and hydroxyapatite (HAp) powders. Adsorption studies were performed to test the effectiveness of several carboxylic acids and alcohols on the dispersion of these powders in the binder system. In both cases, it was found that stearic acid was most effective as a dispersant for the ceramic powder / thermoplastic system. After a suitable dispersant was chosen, ceramic powders were compounded with the binder formulation to yield 55 vol.% ceramic-loaded materials.

After a suitable dispersant was chosen, ceramic powders were compounded with the binder formulation to yield 55 vol.% ceramic-loaded materials. The resultant compound was· used to make filament suitable for use in a modified StratasysTM 3D-Modeler. The filament was well suited for FDC usage, and the parts made using FDC contained no detectable filament-related defects.IntroductionAmong the various techniques that have been used for the fabrication of functional quality ceramic components, Fused Deposition of Ceramics (FDC) has shown great promise. 1 In this process, spooled filament with a diameter of 1.75 mm is fed into a heated liquefier via a set of computer-driven counter-rotating rollers. The liquefier motion is computer controlled in the X-Y plane, and material is extruded through a circular nozzle attached to the end of the liquefier. A foam platform moves in the Z direction, allowing material to be deposited layer by layer.

A foam platform moves in the Z direction, allowing material to be deposited layer by layer. The filament, comprised of ceramic particles in appropriate thermoplastic binders, is used for the fabrication of green ceramic components. When ceramic (or metal) powder-loaded filament is fed into the liquefier, the binder softens and melts inside the liquefier, and acts as a carrier for the particles. The cold filament at the top of the liquefier acts as a piston, creating a positive pressure to extrude the molten material out of the liquefier through the nozzle.

The cold filament at the top of the liquefier acts as a piston, creating a positive pressure to extrude the molten material out of the liquefier through the nozzle. Components are built layer by layer on the platform, and the bonding between neighboring "roads" or beads of material takes place due the "adhesiveness" of the molten binders, combined with the re-melting of previously deposited layers.The motivation for .this work was to develop a family· of thermoplastic •binders suitable for the Fused Deposition (FD) of various ceramics and metals using commercially available thermoplastic polymer constituents. Since the compositions of the various binder components are known, specific binder properties can be tailored based on the requirements of the filler phase and the FDC process. Therefore, a detailed understanding of the various polymer ingredients, and their rolein the final formulation the can be developed.

Therefore, a detailed understanding of the various polymer ingredients, and their rolein the final formulation the can be developed. In addition to compatibility, it is important compatibility of each constituent is critical.Since the new binder will be a multi-component system,613thatsirpilar.·processing exists for.·each component. This includes factors such as melting temperature, decomposition temperature, and decomposition atmosphere. Finally, the binder must exhibit highstrength,high· stiffness, low residue. after burnout, and a relatively low melt viscosity,. particularly when mixed with high solids loadings of ceramic or metaLpowders($O-65vol.%). IThe quality of the feed filament is one of the most important parameters for successful FDC. The. criticaL pr()perti~s for high quality.

IThe quality of the feed filament is one of the most important parameters for successful FDC. The. criticaL pr()perti~s for high quality. filament include viscosity, strength,flexibility, and modulu.s./Jn a~dition to fonning the part once deposited, the filament also acts as a piston at the entrance of the liquefier, forcing molten material out of the nozzle. If the filament exhibits a low stiffness or high viscosity, buckling will occurjust before·.theentrance to the liquefier, thus making FDC impossible. Filaments with high melt vis<;osity.requireeither a very high. stiffness, or a relatively high operating temperature to. lower the melt viscosity to a useable level. Finally, the temperature dependence of. viscosity is also important. A high dependence of viscosity on temperature. allows the filament to maintain consistent mechanical properties up to its entry into the liquefier. This combination ofproperties makes the design of a working filament a.

allows the filament to maintain consistent mechanical properties up to its entry into the liquefier. This combination ofproperties makes the design of a working filament a. challeqge, .Parti<;ularly.forhighly loaded systems.Although a.good binder alone is important, it is not the only requirement for.high solids loading of ceramic powder. The degree of dispersion ofthe powder in the molten thermoplastic binder has a dramatic effect on the rheological properties of resultant FDC filament. In general, the betterthe powder is dispersed in the polymer medium, the lower the resultant viscosity; and by extension, the higher the usable solids loading for a given viscosity range.Studies have been rpade on dispersing ShN4 in thermoplastic binders for FDC applications2 In this cas.e, oleyl alcohol is coated on the powder surface by ball milling in ethanol to provide dispersion for this system. Oleyl alcohoLis a CIS hydrocarbon with an OR functional group.

Since the average size, shape, and surface chemistry varies with the type of powder to be used, there is no guarantee that a surfactant suitable for ShN4 would be suitable in another powder system. Furthermore, the thermoplastic binder formulation used in this work is different from that of the previous ShN4 work.The adsorption behavior of various hydrocarbons onto powder surfaces have been studied for many years3 trends have been established. Ingeneral, polar molecules tend to adsorb onto polar powder surfaces, while non-polar molecules tend to adsorb onto non-polar particle surfaces5 . Oleyl alcohol, like other long-chain alcohols or fatty acids, contains a polar functional group at one end of the chain, and a non-polar hydrocarbon group at these molecules can adsorb onto both non-polar and polar surfaces. A second trend can be seen in the rate at which molecules adsorb into surfaces.

A second trend can be seen in the rate at which molecules adsorb into surfaces. Adsorption from solution is strongly dependent on both the concentration of the adsorbing species, and the chemical nature of the medium in which the adsorption takes place. There are situations where the species of interest remains in solutiotrrather than adsorbing onto the powder surface. This occurs when the species is more compatible with the solvent than the powder surface. The concentration of the adsorbing speciesin.solutionalso playsiacritical role.Based on this earlier work, several,4.the other end.Therefore,Since the powders included in this· study are oxide ceramics, they have a polar surface. Therefore,. we •• can assume that the ..ipolar functional group ofthe surfactant will adsorb onto the powdersurface.

Therefore,. we •• can assume that the ..ipolar functional group ofthe surfactant will adsorb onto the powdersurface. In the case of the surfactantstested (oleic acid, oleyl alcohol, stearic acid, andstearyl alcohol), this would .be the COOHgroup of the acids,or theOHgroup of the alcohols.The ne.w thermoplastic binder formulation developed for the FDC process is comprised entirely of commercially available polymers, <and .. exhibits properties adequate for usage in combination withlead zirconate· titanate (PZT) and. hydroxy.·apatite (HAp) In this paper, the development •of)high-quality ceramic-loaded ceramic powder filler.. feedstock will be outlined-Rheological data from/the raw binder will be presented, in addition to thermal and rheological data from a loaded system. Finally, photographs· of functional components made using this hinder formulation will be presented.ExperimentalFormulations werebatched and prepared using a parts-by-weight measuring system.

Finally, photographs· of functional components made using this hinder formulation will be presented.ExperimentalFormulations werebatched and prepared using a parts-by-weight measuring system. Each formulation consists of atleast one poly-olefin base binder, a hydrocarbon resin tackifier, a polyethylene wax, and a polybutyleneplasticizer.The. various binders, tackifiers,or waxes tested within this system. vary primarily in molecular weight. For convenience,.ithe iterations are labeled ECG-l through ECG..9. The polymers were mixed at 150..175 °c using a high shear mixer, then de..aired until theentrappedairwasiremoved.. Viscosity measurements of the binder formulations were made at 190 °Cusing a Brookfield RVDVTable 1 lists the various fOf1llulations tested within this system.rheometer with a number 27 spindle in conjunction with a Thermocel adapter.The adsorption·of·.four different surfactants was ·measured for the PZT and HAp systems.

For HAp, 50 ml of the solution was mixed with 10.0 g of ceramic powder (surface area = 3.40 m2/g). After mixing, the slurry was centrifuged for 90 min at 1000 rpm to extract the supematent. TGA analysis was then performed on the supematents to determine the concentration of surfactant still in solution after milling and centrifugation. Of the four.surfactatlts, the one with the least surfactant remaining in the supematent was chosen for further studies.Once the suitable surfactant was chosen, the coated powder was mixed with the optimized binder, and rheological measurements were made.. For this portion of the work:, 300.0 g of PZT powder was milled for 4 hours with 300.0 ml of the surfactant solution. The slurry was· then vacuum filtered to remove the solvent and any excess surfactant. After drying, 245.42 g of the coated powder was mixed. with 19.27 g of binder using a Haake System 9000 torque rheometer in conjunction with a 70 cc mixing bowl..

After drying, 245.42 g of the coated powder was mixed. with 19.27 g of binder using a Haake System 9000 torque rheometer in conjunction with a 70 cc mixing bowl.. This corresponds to 55 vol% solids loading. The powder was added in 4 steps spaced 10 minutes· apart, with the initial mixing· temperature set at 160°C. High shear thermoplastic mixing was performed at 140 °c for 20 minutes. Hydroxyapatite / compound was produced in a similar fashion. The viscosity of the PZT / thermoplastic compound was measured using a capillary rheometer. The capillary measured 1.016 mm Shear rates of 2.3 to 230 sec·l were in inside diameter, and· 20.32 mm in length. measured. Measurements were made at 160, 175, and 190 °c.The compounded mixtures were test-extruded at 1 mm / minute and a temperature range of 80 to 95°C into filament 1.75 mm in diameter, cut tolengths ofapproximately In the case of PZT, the purpose was to test the ability to FDC the compounded 50 cm. mixture.

mixture. this filament would be used as the actual feedstock for part production. Once the filament parameters were optimized, continuous lengths of PZT filament were produced using the Haake System 9000 torque rheometer in conjunction with a single screw extruder. Approximately 100 meters of continuous filament wasFor HAp,61640extruded and spooled. This filament was used as feedstock to create green ceramic parts with an upgraded StratasysTM 3-D Modeler™ machine.Fused deposition was performed with both materials using a liquefier temperature setpoint of 140-146 °c, and a build envelope temperature of 25-35 °c. Parts were made using both 250 and 300 J..tm nozzles, depending on part specifications. Build time was 1 to 4 hours, depending on the design.

Parts were made using both 250 and 300 J..tm nozzles, depending on part specifications. Build time was 1 to 4 hours, depending on the design. Compositions with 55 vol.% ceramic were used to make the trial parts.Binder burnout (BBO) was performed in both types of samples using a ramp of In each case, BBO was performed in10.0 °C/hr to 550°C, with a 4.0 hr soak at 550°C. aIr. After BBO, the materials were sintered at 1285 °c for PZT, and 1100 °c for HAp.Results and DiscussionFigure 1 shows the viscosity vs. shear rate plots for the binder formulations tested. In each case, the viscosity of the formulations are relatively low, with ECG-9 showing a viscosity approximately 4 times that of ECG-l. Based on viscosity alone, one may Instead, ECG-9 conclude that ECG-l would make the best binder for FDC applications. was chosen as the binder to be used in future FDC work, primarily due to other factors such as strength and flexibility when mixed in a loaded-system.

was chosen as the binder to be used in future FDC work, primarily due to other factors such as strength and flexibility when mixed in a loaded-system. This point was covered in greater detail in an earlier publication. 6 Figure 2 shows the viscosity vs. shear rate for1000.1 190°C 2 175°C 3 160°C1+ - - - - - - - . . . . - - - - - - - - , - - - - - - - - - -110100100010000Shear Rate (l/see)Figure 2: Viscosity vs. shear rate for 55 vol. % PZT / ECG-6 compound55 vol.% PZT / ECG-6 compounded batch. As expected, the material-is pseudoplastic, with the viscosity of the mix remaining below 1000 Pa-s throughout the range of shear rates tested.Figure 3 shows the thermogravimetric analysis of the supernatant solutions after being mixed with PZT powder for 4 hours. Clearly, stearic acid adsorbed the most onto the powder surface after being mixed with the -PZT powder. A total of 8.1 mg/m2 of617stearic acid adsorbed onto the powder surface from the solution.

The smaller first spike, occuring at approximately 360°C causes open channels to form in the otherwise dense green body. These open channels facilitate removal of the remainder of the binder at approximately 390°C, without blistering or cracking which could present an obstacle to the successful BBO of complex ceramic bodies.Using the new binder formulation along with PZT and HAp powder, multiple copies of complex designs have been produced using FDC. Figure 5 shows photographs of two such designs. Figure 5a is, a photograph of a PZT ceramic part slated for use in Naval towed arrays. '., Figure 5b is a photograph of a green and sintered PZT design also In each case, the ceramic phase was sintered to near for use in Naval towed arrays. theoretical densitY,and was free of defects resulting from FDC processing.

theoretical densitY,and was free of defects resulting from FDC processing. In the case of PZT samples, total firing shrinkage averages 18 % in the radial direction, and 24 % in the z-direction.618..-- .~' ..~~!1 ~~••q:••••••••••..•....••••••. ..••]~:..~~ .• ~£lJ~-~. ;.;.............,,,......"....<~.•••••••••••••• •••••••••••••••··tiI..,•••••••......' •••• ' ......••••••.... ....•••••••••....~.:. .,~a,Oa" '...,...,;.••••....••••'i).••..•••••..••• ••••••••••••......,....."..••••••i···~~·~I;.I··~••••.....'...••••I\<'~•S"-i~:• ..~~l".... ,!, •••••••......""~,,t.,...............

,!, •••••••......""~,,t.,............... ...';'Figure 4: Thermogravimetric analysis of 55 vol.% PZT / thermoplastic compoundFor HAp samples, shrinkage averages 13 % in the radial direction, and 14 % in the z direction.ConclusionsA new binder formulation has been developed for use in Fused Deposition of Ceramics (FDC). This binder formulation consists entirely of commercially available polymers, and exhibits rheological and thermal properties suitable for FDC applications. In addition, stearic acid has been found to provide adequate dispersion for both a lead zirconate titianate-based system and a hydroxyapatite-based system. In both cases, functional ceramic prototypes were manufactured using a Stratasys 3-D Modeler™. Parts made using this new polymer were of high quality, and showed no signs of binder related flaws after sintering.AcknowledgmentsThis work was made possible by ONR ASSERT grant NO.N00014-94-0724. Theauthors would also like to thank D.

Theauthors would also like to thank D. Patel, S. Ocasio, A. Michaels, N. Venkataraman, Bianchi, S. Lee, T. McCarthy, and G.Lous whose contributions made this work possible.619abFigure 5: a) Photograph ofsintered PZT part made via FDCb) Photograph of green and sintered PZT parts made via FDCReferences1. S. Rangarajan, G. Qi, A. Bandyopadhyay, C. Dai, J. W. Han, P. Bharagava, S. Wu, A. Safari, and S. C. Danforth,"The Role of Materials Processing Variables in FDC Process," Proceedings of the Solid Freeform Fabrication Symposium, Edited by H. L. Marcus, J. J. Beaman, J. W. Barlow, D. Bourell, andR. Crawford, The University of Texas at Austin, Austin, Texas, pp. 431 ..440, 1997.2. G. Qi, C. Dai, S. Rangarajan, S. Wu, A. Bandyopadhyay, N. Langarana, A. Safari and S. C. Danforth, "Properties ofRU955 Si3N4 Filament for Fused Deposition of Ceramics, "Proceedings of the Solid Freeform Fabrication Symposium, Edited by L. Marcus, J. . Beaman, J. W. Barlow, D. L. Bourell and R.

The 3S system is able to fabricate intrinsic features without supporting structures; while its downsides exhibit consuming lot of time (30 sec/layer) for fabrication compared to other DLP apparatuses and low efficiency raw material consumption. A new system named as Circular - 3S (C3S) is developed by adapting the 3S technology to improve the fabrication process. It consists of multiple DLP and a circular platform where a paving blade paves the slurry in a circular manner. The demonstrated system has increased the production rate to 200% with printing speed of 15sec/layer. In this paper, the development of the C3S system is presented by simultaneously displaying the capabilities and raw material efficiency of the new C3S system.Keywords: Additive manufacturing, 3D printing, ceramics, slurry, 3S, ceramics, alumina1. IntroductionManufacturing and prototyping of ceramic components in a small-scale can increase the expenses for the molds and processing.

IntroductionManufacturing and prototyping of ceramic components in a small-scale can increase the expenses for the molds and processing. Particularly for the new and complex designs, it is difficult to adapt to produce in a smaller volume at an economical price. Additive manufacturing (AM) can cut down the processing expenses by reducing the number of steps down to only a few. AM allows for the production of structures with complex topologies, while also allowing for the usage of different materials, including metals and alloys (Frazier, 2014), ceramics (Wang, Dommati and Hsieh, 2019), polymers and plastics (Ligon et al., 2017; Dizon et al., 2018).AM with ceramics has existed since the year 1986 (Lee et al., 1986), in which a photopolymerized tape casting process was described. In addition to tape casting, laser scanning and digital light projection are also very common methods of fabricating ceramic products.

In addition to tape casting, laser scanning and digital light projection are also very common methods of fabricating ceramic products. In recent times, there have been major developments in building High Performance Ceramic (HPC) materials with intrinsic features while preserving the thermal, mechanical and corrosive resistance properties of the structure materials. AM systems are now able to build functional grade dense ceramic materials in small scale. (Tang, Chiu and Yen, 2011; Liu et al., 2013; Wang, 2013; Schwentenwein and Homa, 2015; He et al., 2018; Hu et al., 2018).

(Tang, Chiu and Yen, 2011; Liu et al., 2013; Wang, 2013; Schwentenwein and Homa, 2015; He et al., 2018; Hu et al., 2018). Although most of the systems are capable of fabricating the functional grade HPC objects, they tend to exhibit low efficiency in the overall process and are limited to small volumes of parts.1720The Lithography based Ceramic Manufacturing (LCM) based process by Lithoz GmbH (Schwentenwein and Homa, 2015), is a commercial system claims to fabricate 99.3% dense Alumina (Al) objects with 427MPa at 30 μm resolution and maximum volume of 76.8 × 43.2 mm in length and width. The CeraFab 7500 system claims to print up to 100 layers/ hour, i.e. 37.5 sec/layer, and the CeraFab S65 system claims to print up to 150 layers/hour, i.e. 24 sec/layer.

The CeraFab 7500 system claims to print up to 100 layers/ hour, i.e. 37.5 sec/layer, and the CeraFab S65 system claims to print up to 150 layers/hour, i.e. 24 sec/layer. This enables users to fabricate an 8mm thick object with a layer resolution of 25 μm within 3.2 hours and 2 hours respectively.The 3S system developed to fabricate HPC materials is able to fabricate tiny features as small as several hundred microns (Wang, 2013; Wang and Dommati, 2018). However, the 3S system has a build volume of only 57.6 × 32.4 mm in length and width. The ratio for the build volume to the amount of slurry utilized in the 3S system is extremely high which leads to low efficiency of the system. The printing speed of the 3S system is 30 sec/layer.Within industrial applications, it is important to optimize the fabrication time to reduce the wait time and increase productivity.

The printing speed of the 3S system is 30 sec/layer.Within industrial applications, it is important to optimize the fabrication time to reduce the wait time and increase productivity. The first 3S system has been improved to a new process called the Circular 3S system (C3S), in which the platform has a circular build plate and the cycle time for each layer is cut down to half the previous time. By adding another DLP system, the productivity has been multiplied to twice that of the single DLP system. In this paper, a new C3S system apparatus is built and compared to the previous 3S systems and other commercial systems available on the market. The samples 3D printed using C3S system are subjected to sintering and characterized by their flexural strength, density, surface roughness, and shrinkage rate.2. Circular Solvent based Slurry Stereolithography (C3S) systemThe C3S system follows the same working principle as the 3S system.

Circular Solvent based Slurry Stereolithography (C3S) systemThe C3S system follows the same working principle as the 3S system. The slurry is deposited on the build platform, the paving blade paves the slurry uniformly in a circular manner and the DLP system projects the blue light to cure the deposited layer. These steps are repeated until the green parts are formed. The obtained green parts are cleaned under the running water to remove the uncured residual slurry as a part of post processing. The post-processed objects are then sintered at up to 1600 (cid:599) to form completely dense ceramic objects.2.1. Slurry formulationThe slurry is formulated using structure material, solvent, photo-curable resin, dispersant, and dye. The solvent used in this process is methanol while alumina is used as a structure material. The solvent is used to increase the viscosity of the slurry in order to enable homogenous distribution of alumina particles.

The solvent is used to increase the viscosity of the slurry in order to enable homogenous distribution of alumina particles. An orange dye is used as a coloring agent to control the light rays’ penetration during the curing process.2.2. C3S apparatusThe C3S apparatus mainly consists of the following subsystems:1) DLP system 2) Build platform 3) Rotating paving blade17214) Electronics control system 5) Human to Machine interfaceThe whole system is built into a closed cabinet with continuous air circulation and filters to avoid the surrounding dust. The layout of the C3S system is shown in Figure 1.DLP system2-axis systemSlurry depositing systemFigure 1: Layout of C3S system2.2.1. DLP systemThe C3S apparatus is able to host more than one DLP system in order to increase the productivity. In this present apparatus, two DLP systems are installed. However, the current setup is built to enable up to six projectors.

In this present apparatus, two DLP systems are installed. However, the current setup is built to enable up to six projectors. Meaning, the productivity can be directly increased up to 6 times. Six different build volumes can be utilized to fabricate green parts. Figure 2 shows the NVM UV engine used in this process.Figure 2: NVM UV EngineThe NVM UV engine is 2K resolution, naked body type setup that is compact in size. Hence, it is easy to accommodate multiple of such DLP systems.2.2.2. Build platform1722The build platform in this apparatus is a stainless steel square cut non-erosive plate on which the slurry is paved in a circular ring. The circular shape enables the slurry to pave over the surface continuously without letting it go waste, which is unlike the old 3S system where the excess slurry that is deposited for each layer goes waste. The build platform is setup with a linear actuator to carry it decremented or incremented in the Z-axis.

The build platform is setup with a linear actuator to carry it decremented or incremented in the Z-axis. Figure 3 shows the layout of the circular platform with up to 6 DLP configurations. In this paper, the P2 and P5 areas are used for exposing masking light on the slurry ring with two DLP systems.02000100Scafp,l"r,_,=';::':::➔--.J.L_._ ScraperbladeBuild platf'onnCircular ring Ullldc'.:.---4-..,_from tht> dcpositc<I slurryFigure 3: Build platform with circular paving2.2.3. Rotating paving bladeThe paving blade is used to pave the slurry into the circular ring over the build platform. It rotates continuously at a controlled speed. The paving blade design is as shown in the figure 4 (a), while Figure 4 (b) represents the setup of the paving blade in the apparatus.Ball setArm0p~,Al. ~wn00000Ball se1Paving blade(a)(b)Figure 4: (a) Paving blade design (b) Paving blade setup2.2.4.

~wn00000Ball se1Paving blade(a)(b)Figure 4: (a) Paving blade design (b) Paving blade setup2.2.4. Electronics control systemThe C3S apparatus consists of a slurry dispensing system, a stirring system within the slurry tank, and an electronic control system that controls the whole setup.1723The slurry dispensing system collects the slurry from the slurry tank through a pump and disperses it over the build platform. The pump motor speed is controlled in open loop and the input pulses are fixed to maintain the equal proportion of the slurry to be dispensed for every layer. The stirring system in the slurry tank is required to keep the slurry homogenousIn this setup, an open source Arduino board is used as the motherboard to control the Z-axis motion, paving blade motor control, masking light ON/OFF control, stirring motor ON/OFF, and the slurry dispensing pump.

Figure 5 (a) and (b) shows the stirring system and the electronics control system.Stirring motorCouplerPaving blade speed controlCovering platepum mg SI~ 1peElectronics LayoutPower ON/OFFSlurry tankStirring blade(a)(b)Figure 5: (a) Stirring system (b) Electronics layout2.2.5. Human Machine InterfaceAn interface is developed to control the apparatus and feed the geometric codes to the machine from the computer. The interface is integrated with a slicing code that can directly slice the 3D file and send the control commands to the machine. The interface is as shown in the figure 6.3D model3D model layerLayer status110 \ fF.Connect/ disco,mect Z ax.is origin buttonAuto paramctcTS settings Instructions scriptP2Projector selectionmanual control toggle buttonsFigure 6: Interface layout2.3.

Fabrication of green parts17243D models and desired print parameters are fed to the user interface in Figure 6, which sends the information to the machine, which fabricates green parts layer-by-layer. The layer resolution that is possible in this apparatus is between the ranges of 10 – 50 μm range. With a recommended layer resolution of 20 μm and pixel resolution of 30 μm, the green parts can be fabricated at the rate of 15 sec/layer. In C3S system, the paving blade is not required to go backwards to initiate paving the new layer; this cuts down the drying time starting from the first cycle of slurry paving. Unlike the linear 3S system, where the paving blade needs to wait until the paved layer dries and returns to its origin.Figure 7 (a), (b) & (c) represents the flow chart of the fabrication of green parts. Initially, the slurry is deposited on the build platform and a paving blade paves the slurry in the form of a ring as shown in 7(a).

Initially, the slurry is deposited on the build platform and a paving blade paves the slurry in the form of a ring as shown in 7(a). Next, the masking light is projected by the DLP systems, which is P2 and P5 in this case. The masking light is exposed for 0.7 sec each layer for curing. Finally, the build platform is removed from the system and the green parts are washed with water until the uncured slurry is gone.Figure 7: (a) Paving of the deposited slurry on the build platform (b) P2 and P5 projectorsexposing the masking light (c) Post processing the green parts under running waterTable 1 shows the data of the dimensions of the green parts obtained after 3D printing. The design sample is 6mm x 6mm x 1.8mm. The green parts obtained from the experiment with 15 sec/layer printing speed exhibited bigger dimensions than the parts obtained from 30 sec/layer printing speed experiment.

The green parts obtained from the experiment with 15 sec/layer printing speed exhibited bigger dimensions than the parts obtained from 30 sec/layer printing speed experiment. It is due to the vaporization of the solvent from a paved layer, causes increase in vol % of the dye. At low printing speeds the vaporized solvent is high at high printing speeds. The increase in vol % of dye decreases the penetration of light through the paved layer. Hence, at the same exposure time of masking light, the parts printed at 15 sec/layer penetrate more light energy than parts printed at 30 sec/layer.

Hence, at the same exposure time of masking light, the parts printed at 15 sec/layer penetrate more light energy than parts printed at 30 sec/layer. More light energy can cause over curing and more diffuse scattering, which results in increase in the dimensional errors.Table 1: Dimensions of the green partsSampleLength (mm)15 sec/layer Width (mm)Thickness (mm)Length (mm)30 sec/layer Width (mm)Thickness (mm)1 2 3 46.24 6.07 6.18 6.066.24 6.12 6.11 6.111.93 1.97 1.92 1.946.03 5.95 6.00 6.036.09 5.95 6.02 6.021.86 1.83 1.87 1.8617255 6 7 86.13 6.10 6.10 6.126.11 6.11 6.07 6.101.94 1.94 1.92 1.946.02 5.90 5.87 6.105.94 5.93 5.90 6.041.83 1.81 1.85 1.87Average Deviation6.125 0.0526.121 0.0441.938 0.0145.988 0.0675.986 0.0581.848 0.0192.4. Debinding and sintering of the green partsDebinding is the process of removing the binder from the green parts.

Debinding and sintering of the green partsDebinding is the process of removing the binder from the green parts. The debinding is done by subjecting the green parts to heating up to 350 (cid:599) at the rate of 3.4 (cid:599)/min for 1 h 30 min and then from 350 (cid:599) to 600 (cid:599) at the rate of 0.92 (cid:599)/min for 4 h 30 min. This is followed by a sintering process, which is done by heating the parts up to 1600 (cid:599) at the rate of 1.6 (cid:599)/min. At 1600 (cid:599), the green parts become highly densified and exhibit pure ceramic properties. The parts are left at 1600 (cid:599) for two hours to let the densification process complete. After densification, the objects are cooled down to 40 (cid:599) at the rate of 2.8 (cid:599)/min for 9 hours.

From the data, it is observed that the dimensions of the low speed printed parts are less than the parts printed at high speed.Table 2: Dimensional data of the sintered partsSample15 sec/layer (high speed) Width (mm)Length (mm)Thickness (mm)30 sec/layer (low speed) Width (mm)Length (mm)Thickness (mm)14.864.861.584.824.851.4524.854.821.544.924.891.5134.834.891.514.84.841.4744.874.91.474.864.871.49172654.844.851.504.774.781.5264.844.881.544.824.861.574.814.831.594.774.841.518 Average Deviation4.834.8410.0174.814.8550.0291.521.5310.0354.844.8250.0444.824.8440.0291.531.4980.0233. Characterization of the samples printed at different layer cycle timeIn this paper, the mechanical properties are characterized by printing the sample parts at different speeds with the same slurry material.

The printing speeds are 15 seconds/ layer (high speed) and 30 seconds/layer (low speed). 3D model of blocks with 6mm in length and width and thickness of 1.8 mm are 3D printed to determine the shrinkage rate, average volume and density. To determine the flexural strength, rectangular bars with dimensions 30×2.4×1.8 mm in length, width, and thickness respectively, are fabricated. The printed green parts and the sintered parts of 3D models are as shown in the figure 9 (a) and (b).~J...£"i,... ' " • '2C3S 15 sec/ layerI,.'.t"I...3f,51 0.. ~ ,.\.' l'IC3 S 30 sec/ I ayer... ,,. .... .\tI4.' =t • l 'C3 S 15 seconds" ... "' • , . s,.1.·• .... • .....~Ii., r~ l"f ·-" 1,i' ' 'IC3S 30 seconds(a)(b)Figure 9: (a) Green parts (b) Sintered parts3.1. Shrinkage rateThe dimensions of the obtained green parts are measured under the 3D macroscope.

Shrinkage rateThe dimensions of the obtained green parts are measured under the 3D macroscope. The table 1 represents the length, width and thickness of the green parts obtained from C3S printing with 15 sec/layer and 30 sec/layer. From the table 3 the shrinkage rate of the parts printed at 30 sec/layer speed is slightly lesser than the parts printed at 15 sec/layer speed. The shrinkage is almost linear in the X, Y and Z directions with slight error in length of the parts printed at low speed. The mean shrinkage rate of the parts printed at 15 sec/layer is 20.95 × 20.68 × 20.96 % and the parts printed at 30 sec/layer is 19.41 × 19.08 × 18.93 % in X, Y and Z directions respectively.

Comparatively, the low speed printed parts exhibited the low shrinkage rate due to the occurrence of high solid loading with increased amount of solvent vaporization every layer while printing.Table 3: Shrinkage rate of the samples after sinteringSample15 sec/layer30 sec/layer1727Length (%)Width (%)Thickness (%)Length (%)Width (%)Thickness (%)1 2 3 4 5 6 7 8 Average Deviation22.115 20.099 21.845 19.637 21.044 20.656 21.148 21.078 20.953 0.72722.115 21.242 19.967 19.804 20.622 20.131 20.428 21.148 20.682 0.68618.135 21.827 21.354 24.227 22.680 20.619 17.188 21.649 20.960 2.03720.066 17.311 20.000 19.403 20.764 18.305 18.739 20.656 19.406 1.06720.361 17.815 19.601 19.103 19.529 18.044 17.966 20.199 19.077 0.89922.043 17.486 21.390 19.892 16.940 17.127 18.378 18.182 18.930 1.7273.2. DensityThe density of the samples is measured using a densimeter that follows Archimedes principle.

DensityThe density of the samples is measured using a densimeter that follows Archimedes principle. The obtained mean density value of the eight samples is measured as 3.846 g/cm3 for 15 sec/layer and 3.874 g/cm3 for 30 sec/layer speed. From the data presented in table 4, the density of the low speed printed parts is 0.028 g/cm3 higher than the high speed printed parts. This variation is due to the increase in solid loading every layer. At low speed printing, the vaporization of solvent is high causing the solid particles to come closer and hence fixated at the masking light with a higher density. Whereas at high-speed printing, the presence of solvent is relatively high and caused the solid loading little lesser. The difference in standard deviations for high speed and low speed printing is also very high. The inconsistency of density in high-speed printed parts is due to presence of internal cracks or pores in few of the samples.

The inconsistency of density in high-speed printed parts is due to presence of internal cracks or pores in few of the samples. Whereas in low speed printed parts the density values are consistent causing lower standard deviation.Table 4: Density of the samples after sinteringSample15 sec/layer30 sec/layer1 2 3 4 5 6 7 8 Average Deviation3.67 3.99 4.04 3.83 3.64 3.98 3.96 3.66 3.846 0.1571573.86 3.79 3.96 3.81 3.91 3.84 3.96 3.86 3.874 0.0599873.3. Flexural strength1728The flexural strength of the workpieces is measured by the 3-point bending test as shown in figure 10, which is based on the ASTM C1161 standards. In order to obtain the final 3D printed parts with the dimensions of 25×2×1.5 mm length, width and depth respectively, the shrinkage rate is predicted to be 20% and hence the 3D model with 30×2.4×1.8 mm in length, width and depth is 3D printed and sintered. The obtained final parts are in the tolerances mentioned as per the ASTM C1161 standards.

The obtained final parts are in the tolerances mentioned as per the ASTM C1161 standards. The mean flexural strength obtained for the eight rectangular bars is measured to be 367.72 MPa, 418.29 MPa for parts printed at 15 sec/layer and 30 sec/layer speeds. The table 5 shows the results of the 3-point bending test on the printed samples. The difference in the flexural strengths is due to the higher density of the obtained samples at low speed printing.Figure 10: 3-point bending testTable 5: Characterization of the sintered samples15 sec/layer 339.11 1 339.49 2 451.44 3 300.64 4 378.42 5 423.97 6 396.99 7 311.73 8 Average 367.72 Deviation 47.60Sample30 sec/layer 455.51 432.46 335.03 415.17 319.89 373.06 461.58 489.82 410.32 54.514. Discussion1729The C3S system is operated at two different printing speeds i.e. at 15 sec/layer and 30 sec/layer. At high printing speed, the fabrication of the green parts is reduced to half the time it usually takes.

at 15 sec/layer and 30 sec/layer. At high printing speed, the fabrication of the green parts is reduced to half the time it usually takes. However, the mechanical characteristics for the parts printed at low speed exhibited better results compared to the parts printed at high speed. Due to the high vaporization of solvent in low speed printed parts, it resulted in high solid loading and higher density of the parts, 3.87 g/cm3 whereas it is 3.84 g/cm3 at higher speed. In addition, the flexural strength of the parts printed at low speed is 410.32 MPa, which is ~42 MPa higher than the parts printed at high speed. For a 6×6×1.8 mm 3D model, after printing at high and lower speeds, the dimensional error in the green parts is measured to be comparatively lesser in the low speed printed parts. It is due to the raise in dye vol % when the solvent is vaporized at a proportional rate of the paving time for every layer. The masking light scattering is inversely proportional to the vol % of dye.

The masking light scattering is inversely proportional to the vol % of dye. Hence, the dimensional error that is caused due to overcuring is lesser in low speed parts. To obtain green parts with exact dimensions of the 3D model, the printing speed time can be optimized. Although it is impractical to achieve parts with exact dimensions as 3D model, the parts can be printed within the standard tolerances.5. ConclusionA new C3S system is developed based on the same working principle as 3S to increase the productivity. The new apparatus is suitable for small-scale and batch wise production processes. By implementing two projectors in the C3S, the productivity has been increased 4 times; counting the printing speed when reduced to half. The new C3S is capable of printing at different speeds effectively. The mean shrinkage rate of the parts is 20.95 × 20.68 × 20.96 % and 19.41 × 19.08 × 18.93 % in X, Y and Z directions respectively for the parts printed at 15 sec/layer and 30 sec/layer.

The mean shrinkage rate of the parts is 20.95 × 20.68 × 20.96 % and 19.41 × 19.08 × 18.93 % in X, Y and Z directions respectively for the parts printed at 15 sec/layer and 30 sec/layer. The parts printed at 15 sec/layer speed exhibited flexural strength of 367 MPa, 3.84 g/cm3 density; parts printed at 30 sec/layer speed exhibited flexural strength of 418 MPa, density of 3.87 g/cm3. The layer resolution remains the same while the mechanical properties depends upon the chosen printing speed. To increase the productivity and make an AM system that is suitable for batch production, the C3S can be utilized to its full potential by implementing up to six projectors in one system.References:Dizon, J. R. C. et al. (2018) ‘Mechanical characterization of 3D-printed polymers’, Additive Manufacturing. Elsevier B.V., 20, pp. 44–67. doi: 10.1016/j.addma.2017.12.002.Frazier, W. E. (2014) ‘Metal additive manufacturing: a review’, Journal of Materials Engineering and Performance.

velopment of a Co-Cr-Mo to Tantalum Transition using LENS for Orthopedic ApplicationsSaripella Surya Kumar, IIT Bombay Brent Stucker, Utah State UniversityAbstractBiomedical implant material research using additive manufacturing is a popular field of study. Many potential material combinations exist which, if implemented properly, could have a significantly positive effect on implant life and functionality. One material combination of interest is attaching porous Ta bone ingrowth material to a CoCrMo corrosion and wear resistant bearing surface. An investigation of the ability of the LENS process to join Ta to CoCrMo was undertaken. Direct joining of CoCrMo to Ta was known to be problematic, and thus transitional layers of other biomedically-compatible materials were investigated.

Direct joining of CoCrMo to Ta was known to be problematic, and thus transitional layers of other biomedically-compatible materials were investigated. It was determined that a transitional layer of zirconium appeared to be the best transitional material for this application due to its excellent biocompatibility, followed by stainless steel, with a lesser biocompatibility but better adhesive properties.IntroductionThe development of multi-material and graded material structures is revolutionizing material science and engineering. This is due to the fact that a wide range of properties can be obtained, leading to more optimized properties and the development of new material systems. The material combination for the study considered here is a graded structure from porous Tantalum to Co-Cr-Mo. This material combination is desirable for biomedical implant structures where one side of the implant is a bone in-growth structure and the other side of the implant is a bearing surface.

This material combination is desirable for biomedical implant structures where one side of the implant is a bone in-growth structure and the other side of the implant is a bearing surface. A typical problem encountered when coating porous Tantalum directly with Co-Cr-Mo is the process of crevice corrosion. Thus, a transitional material, which blends well with both Co-Cr-Mo and porous Tantalum, was sought. A number of bio-compatible materials were considered for this application, including Titanium, Zirconium, Rhenium, Niobium and Stainless Steel, as potential candidates for the interfacial layer. An investigation was undertaken to ascertain the best biocompatible transition metal between Co-Cr-Mo and Ta using an Optomec LENS 750 machine.BackgroundAccording to Hubler[1] thin film layers can be used to protect implant materials. These layers lead to an improvement in corrosion resistance. The properties are largely dependent on the adhesive properties and the layer thickness.

These layers lead to an improvement in corrosion resistance. The properties are largely dependent on the adhesive properties and the layer thickness. They also depend on the possibility of formation of intermetallics between the metals being used. The possibility273of intermetallic formation when layering materials necessitates looking to phase diagrams before choosing materials as interfacial layers.Ensinger[2] (et al.) have shown that the implanting of noble metals onto the surface of materials leads to an increase in corrosion resistance. Hence it is beneficial to choose a biocompatible material with low reactive properties to be used as the interfacial material. This would lead to retaining of the material properties of the implant substrate and also form an excellent coating for depositing the strongly corrosion resistant Co-Cr-Mo.

This would lead to retaining of the material properties of the implant substrate and also form an excellent coating for depositing the strongly corrosion resistant Co-Cr-Mo. They have shown that Tantalum suffers Hydrogen embrittlement when it corrodes in strong acids and so the use of materials like Zirconium, Rhenium and Niobium as interfacial layers is encouraged.Matsuno[3] (et al.) have shown that Titanium, Zirconium, Rhenium and Niobium metals are biocompatible and have excellent blending with the tissues in the body. The experiments conducted by them indicate the fact that these materials have excellent corrosion resistance and can act as excellent intermediate surfaces.Trillo [4] (et al.) have shown that Titanium and Tantalum form excellent solid solutions having properties comparable to that of Ti-6Al-4V. Also they have reported excellent hardness values.

Also they have reported excellent hardness values. This strengthens the belief that these metals can be used as intermediate layers for implant production.Cui [5] (et al.) have assessed the Rhenium-Tantalum system in great detail. They have indicated that Rhenium has excellent fatigue and creep failure resistance. Also the fact that the phase diagram has no intermetallics strengthens the fact that it will act as an excellent intermediate layer.Experimental EquipmentLaser Engineered Net Shaping (LENS) produces functional metal prototypes that have excellent material properties. LENS also has a unique ability to build parts with novel geometries and multi-material as well as functionally gradient material compositions, opening the door for production of breakthrough end-products. As an additive process, LENS allows users to modify or rework existing prototypes for design iterations, rather than fabricating an entirely new part with each change.

As an additive process, LENS allows users to modify or rework existing prototypes for design iterations, rather than fabricating an entirely new part with each change. LENS has the potential to be used to cost-competitively manufacture surgical devices and anatomically customized, as well as standard-sized, medical implants with a range of functional enhancements that improve wear characteristics and quality of life for the product.The LENS process takes place in a controlled atmosphere processing chamber. The system provides gas recirculation to remove any Oxygen that leaks into the system. The atmosphere inside is inert Argon. The presence of Argon reduces the chances of oxidation within the chamber. The continuous wave, high power Nd:YAG Laser (500- 600W) operates at 1.064µm. The laser is mounted on the top of the system and is delivered to the work surface via a beam delivery system.

The continuous wave, high power Nd:YAG Laser (500- 600W) operates at 1.064µm. The laser is mounted on the top of the system and is delivered to the work surface via a beam delivery system. The energy from the laser is274used in melting the powder and depositing it onto the substrate material. A 3-axis motion-controlled table provides for movement of the part being fabricated.Several parameters affecting the LENS process include Laser Power, Laser Traverse Speed and Powder Feed Rate.

A 3-axis motion-controlled table provides for movement of the part being fabricated.Several parameters affecting the LENS process include Laser Power, Laser Traverse Speed and Powder Feed Rate. These parameters were investigated to determine whether the various materials under consideration were capable of being processed using LENS.Benefits of LENS for use in development of biomedical implants include:Enhancing components by adding layers of wear-resistant coatings, hard-facing, and other surface treatments.Lower per-piece production costs for low-volume or custom production • • Broad range of materials – metals, metal alloys, ceramics, composites • Novel geometries, material compositions and textures, to:Improved material properties for longer part life1. Reduce implant wear at key points by optimizing the placement of alloys 2.

Reduce implant wear at key points by optimizing the placement of alloys 2. Foster implant integration with native tissue by using multiple biocompatible materials that blend with the tissue in the body more optimallyFor the sintering experiments described below, a high temperature furnace is employed. The parameters that can be controlled in the high temperature furnace include the temperature and the pressure in the furnace as functions of the time. Also the atmosphere inside the system is inert (argon) to minimize oxidation.Experimental ProcedureFor the experiments performed, a simple observation of the compatibility of LENS- deposited material combinations was observed by looking for evidence of delamination. For some specimens this involved cutting and polishing the specimens and then observing the interface using optical microscopy.

For some specimens this involved cutting and polishing the specimens and then observing the interface using optical microscopy. In most cases delamination would occur prior to cutting and polishing if the materials were incompatible or the procedure did not result in proper layer adhesion.The materials used for the experiments were Zirconium powder (-100/+250 mesh size) obtained from Zirconium Research, Co-Cr-Mo powder and Stainless Steel (316L) powder obtained from Carpenter Technologies, Porous Tantalum substrates from Zimmer, Tantalum Metal Foil (0.127mm thick) from Alfa Aesar, and Co-Cr-Mo and Ti- 6Al-4V substrates obtained from Medicine Lodge, Inc.

Rhenium and Niobium, although potentially optimal materials, were not investigated due to the high cost of the raw materials.It should be noted that when using a porous Ta substrate, conduction cooling is minimized due to the relatively small conductive paths through the porous structure, as compared to a traditional, solid LENS substrate where conduction cooling occurs rapidly.275Thus, LENS parameters used for solid substrate depositions will not necessarily be optimum for those on porous substrates. In particular, the cooling rates and the resultant microstructures will vary.Although the thickness of the deposits was not recorded precisely, each layer would typically result in a deposit of approximately 0.010”.

Thus, for example, a 3 layer deposit would result in a 30 thousandths thick layer.Experimental Results & DiscussionCo-Cr-Mo to Ta depositionsFrom previous research it was observed that crevice corrosion might lead to problems when directly depositing Co-Cr-Mo onto a porous Tantalum substrate. The direct deposition of Co-Cr-Mo onto Tantalum using LENS provided for excellent adhesion, but the crevice corrosion resistance of these deposits was not investigated for this project.A major drawback of all of the LENS depositions of Co-Cr-Mo was the surface finish. The surface formed is highly uneven and this unevenness increased up as the amount of Co-Cr-Mo increased. This might be due to the surface tension characteristics of Co-Cr- Mo in the molten phase.Ta foil to porous TaThe next step was to observe whether we could laser-deposit Tantalum foil onto the porous substrate.

If a thin layer of solid Ta could be formed at the surface of the porous Ta, then it might be possible to deposit Co-Cr-Mo on top of the solid Ta and eliminate the possibility of crevice corrosion due to the elimination of crevices present at the interface. Laser glazing of the Tantalum foil onto the porous Tantalum substrate was investigated, without full melting of the foil. It was observed that the Tantalum foil did not stick properly to the substrate due to the fact that there was very limited bonding area.A series of experiments testing laser powers which resulted in melting of the Ta foil also occurred. During formation of a melt pool, the foil started forming bead-like structures which led to the segregation of the Tantalum foil rather than a uniform distribution over the substrate.

During formation of a melt pool, the foil started forming bead-like structures which led to the segregation of the Tantalum foil rather than a uniform distribution over the substrate. Warping of the foil was observed due to temperature gradients within the foil, and thus, after numerous tries, the idea of laser welding a Ta foil to the porous Ta was abandoned.Transitional MaterialsFor the development of transitional materials, Titanium, Zirconium and Stainless Steel were chosen. From the various phase diagrams of interest it was observed that Zirconium does not form any intermetallic compounds with Cobalt, Chromium, Tantalum or Molybdenum whereas Titanium formed a few intermetallic compounds with both Cobalt276and Chromium.

Nothing definitive could be predicted about the behavior of Stainless Steel because of the large number of elements within Stainless Steel.Intermetallic compound formation often results in crack propagation due to the brittle nature of the material and the concentration of stresses at the intermetallic due to its thermal expansion and stiffness differences with respect to the matrix metal alloy. Hence it is important to avoid large-scale intermetallic formation if at all possible.Another key material property is the coefficient of thermal expansion. A transitional material should theoretically have a CTE between the materials of interest in order to minimize CTE mismatches.Ti to Co-Cr-MoInitial experiments were performed using Titanium as the interfacial layer. It was observed the Titanium stuck very well to porous Tantalum. But when deposition of Co- Cr-Mo were added to the deposited Titanium, they started to delaminate.

It was observed the Titanium stuck very well to porous Tantalum. But when deposition of Co- Cr-Mo were added to the deposited Titanium, they started to delaminate. This could be either due to the build-up of residual stresses as a result of differences in CTE and the large temperature changes during the process or it might be due to intermetallic formation between the Ti and Co-Cr-Mo, or combinations of the two.An investigation of the influence of parameter changes on depositions of Co-Cr-Mo onto Ti-6Al-4V substrates was performed to determine inherent compatibility problems. For the depositions, the following were changed:1. Laser Power was varied from 33W – 272W 2. Powder feed rate in the range of 0.48-0.60 cc/min and 3. Laser Traverse speed from 10-30 inches/min.In addition, laser scanning pattern was varied.

Laser Power was varied from 33W – 272W 2. Powder feed rate in the range of 0.48-0.60 cc/min and 3. Laser Traverse speed from 10-30 inches/min.In addition, laser scanning pattern was varied. There is a possibility that if the laser scanning pattern is rastered along a single direction from layer-to-layer, the development of residual stresses in that particular direction will lead to delamination. To investigate this possibility, depositions were done such that successive layers were at an angle of 120° with respect to each other. Also, some samples were allowed to cool down after each layer was deposited, which reduced the severity of delamination.It was observed that in almost all cases there was some delamination immediately when the third layer of Co-Cr-Mo was being deposited. In the single case where delamination did not occur in the machine, the layers peeled off when the deposit was sand blasted.

In the single case where delamination did not occur in the machine, the layers peeled off when the deposit was sand blasted. This showed that there was inherently poor bonding between Co-Cr-Mo and Ti-6Al-4V. Hence the search for optimized LENS parameters was abandoned for this material combination.Ti to ZrSince Titanium stuck well to Tantalum, it was proposed to study the characteristics of the deposition of Zirconium onto Titanium. The possibility of a graded implant with both277Titanium and Zirconium as interfacial materials was thusly investigated. An experiment was conducted to ascertain the deposition characteristics of Zirconium on Ti-6Al-4V. The deposits appeared to give excellent adhesion. A set of optimum parameters were:1. Laser Power ~ 153 W 2. Laser Traverse speed = 30inches/min. 3. Powder feed rate ~ 1.13cc/min.Zr to TaNext Zirconium was deposited onto the porous Tantalum substrate.

Laser Power ~ 153 W 2. Laser Traverse speed = 30inches/min. 3. Powder feed rate ~ 1.13cc/min.Zr to TaNext Zirconium was deposited onto the porous Tantalum substrate. From preliminary experiments it was quite clear that Zirconium stuck very well to Tantalum. After many iterations, excellent deposits of 0.9inch diameter were created on porous Ta specimens. Each layer of the deposit was at an angle of 60° clockwise with respect to the previous layer. The number of layers in each deposit was six. After depositing each layer the sample was allowed to cool for about 15-20 min. so that the residual stresses (if any) would be relaxed and also the chances of warping of the sample at high temperatures were reduced greatly. The LENS machine parameters worked out well and the deposits came out in excellent fashion. Once the depositions were completed onto the porous Tantalum pieces the layers were machined to obtain a flat layer of Zirconium on the substrate.

Once the depositions were completed onto the porous Tantalum pieces the layers were machined to obtain a flat layer of Zirconium on the substrate. The LENS parameters for the deposition process were as follows:1. Laser Power ~ 153W 2. Powder feed rate ~ 1.45cc/min. 3. Laser Traverse speed = 30 in/min.Zr to Co-Cr-MoThe deposition of Co-Cr-Mo onto Zirconium was also studied. For initial studies, Zirconium powder was deposited onto a Mild Steel plate (so as not to deplete the supply of porous Ta specimens) followed by depositions of Co-Cr-Mo onto the deposited Zirconium. The deposits were of rectangular profile. During these deposits it was observed that Co-Cr-Mo delaminated slightly during deposition of 6 layers, even after significant optimization of LENS parameters. However, the bonding was significantly better than that of Ti and Co-Cr-Mo. One of the important similarities in all the delaminations is that the deposits started delaminating from the edge.

However, the bonding was significantly better than that of Ti and Co-Cr-Mo. One of the important similarities in all the delaminations is that the deposits started delaminating from the edge. This strengthens the belief that the delaminations were a result of CTE mismatch at the material interface.It was determined that deposits three layers thick stuck quite well. These layers did not peel off even after sand blasting. This is apparently due to a lesser build-up of residual stresses after 3 layers than after 6 layers. The deposits of Co-Cr-Mo onto Zirconium that bonded well had process parameters as follows:1. Laser Power ~ 93W 2. Powder Feed Rate ~ 0.81cc/min. 3. Laser Traverse Speed = 30 in/min.278One rectangular deposit of Zirconium onto solid Co-Cr-Mo was carried out. This was to see whether delamination occurred in this case. As expected, due to CTE mismatch, the layers delaminated after only two layers of deposition. The parameters for the study were:1.

This was to see whether delamination occurred in this case. As expected, due to CTE mismatch, the layers delaminated after only two layers of deposition. The parameters for the study were:1. Laser Power ~ 93W 2. Powder feed rate ~ 1.13cc/min. 3. Laser Traverse Speed = 30 in/min.Solid State SinteringAfter these depositions were carried out an attempt was made to utilize solid-state sintering to join a Co-Cr-Mo piece to the LENS-deposited Zirconium layer on the porous Tantalum. A temperature of 1250°C was used for this experiment. The process was carried out in an inert atmosphere of Argon so that any chances of oxidation due to the presence of atmospheric Oxygen would be minimized.

The process was carried out in an inert atmosphere of Argon so that any chances of oxidation due to the presence of atmospheric Oxygen would be minimized. The process was controlled with the temperatures and pressures varying with time as shown in the graphs.Variation of Pressure with time in High Temperature Furnace0, 012.5, 20017.5, 20029.5, 7600100200300400500600700800051015202530Time (hrs)Pressure in chamber (mm of Hg)279Variation of Temperature with Time for the High Temperature Furnace0, 2712.5, 125017.5, 125029.5, 270200400600800100012001400051015202530Time (hrs)Temperature inside Chamber (C)The result of the sintering experiment showed mixed success. The apparatus we used did not enable applying pressure to aid the bonding process, and only gravitation force was applied. Even so, limited adhesion did occur.

The apparatus we used did not enable applying pressure to aid the bonding process, and only gravitation force was applied. Even so, limited adhesion did occur. It appears that adequate solid state sintering may be possible under more carefully controlled conditions at higher pressures.Stainless Steel to Co-Cr-Mo, Zr and TaA few experiments were also conducted using Stainless Steel as an intermediate material. First, the Stainless Steel(316L) was deposited onto a Mild Steel plate and then Co-Cr-Mo onto the Stainless Steel. We observed that there was excellent sticking and no signs of delamination were visible. The parameters for these deposits are:1. Laser Power ~ 93W 2. Laser Traverse Speed = 30in/min. 3. Powder Feed Rate in the range of 0.81-1.67cc/min (all work fine)It was also noted that Stainless Steel stuck very well to both Zirconium as well as Tantalum. For these deposits: 1. Laser Power ~110W 2. Laser Traverse Speed = 30in/min. 3.

Adhesion at 3 layers and below, but delamination otherwise. Excellent adhesion Excellent adhesionConclusionsIt was determined that there are several possible methods for creating a Co-Cr-Mo to Ta transition using LENS. Although Stainless Steel is not considered an ideal material for long-term implant use, it has fair biocompatibility and good adhesion as a transitional material. Zirconium appears to be the best biocompatible transitional material. LENS of Zr to porous Ta proceeds smoothly, and additional LENS deposits of Co-Cr-Mo can be utilized if only a thin coating (less than 3 layers to minimize residual stresses) is required. An additional manufacturing plan might utilize LENS to deposit Zr and then solid state sintering under pressure to join the Zr to a thicker Co-Cr-Mo layer.

An additional manufacturing plan might utilize LENS to deposit Zr and then solid state sintering under pressure to join the Zr to a thicker Co-Cr-Mo layer. Although Rhenium and Niobium were not studied, due to financial constraints, these materials would likely work well for this applications, as their material characteristics match well the needs of a good transition material between Ta and Co-Cr-Mo.References[1] R.Hubler: in ‘Characterisation of gradient interfaces in thin film multilayer used to protect orthopaedic implants`, Surf. Coat. Technol. ,p 116-119,1999.[2] W.Ensinger, G.K.Wolf.: in ‘Electrochemical and Corrosion Protection properties of Ion implanted thin films’, Electrochimica Acta, Vol. 39, No. 8/9, pp.1159-1164, 1994.[3] H.Matsuno, A.Yokoyama, F.Watari, M.Uo, T.Kawasaki. : in ‘Biocompatibility and Osteogenesis of refractory implants, Titanium, Hafnium, Niobium, Tantalum and Rhenium`, Biomaterials, Vol.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDEVELOPMENT OF A CUSTOMIZED CPAP MASK USING REVERSE ENGINEERING AND ADDITIVE MANUFACTURINGZhichao Ma*, Javier Munguia, Philip Hyde, Michael DrinnanSchool of Engineering. Newcastle University, UKAbstractContinuous positive airway pressure (CPAP) therapy has been widely used to treat moderate and/or severe Obstructive Sleep Apnoea (OSA) syndrome since its invention. However, CPAP mask interface induced side effects, such as air leak, noise, discomfort and facial skin problem, considerably affect the overall effectiveness of CPAP treatment. Conventional CPAP masks designed with averaged individual facial characteristics have standard configuration and limited material selection. Mask size ranges are limited, only coming with small, medium, and large.

Mask size ranges are limited, only coming with small, medium, and large. The material used for the mask interface fabrication is mainly silicone-based material. Besides the limitations on mask configuration and material, there are no comprehensive mask selection templates and guidance offered by CPAP mask suppliers. Individuals have completely different physical characteristics, such as face topology, skin sensitivity, the severity of OSA syndrome, sleep habit and breathing pattern. Therefore, conventional masks cannot properly fit individual’s physical characteristics. Customization of CPAP mask using Reverse Engineering and Additive Manufacturing techniques offers the great potential to minimize the CPAP mask interface induced side effects.1. IntroductionObstructive Sleep Apnoea (OSA) syndrome is a typical type of sleep-disordered breathing syndrome. Continuous Positive Airway Pressure (CPAP) therapy has been widely used to treat OSA syndrome for several decades.

Continuous Positive Airway Pressure (CPAP) therapy has been widely used to treat OSA syndrome for several decades. However, compliance and dissatisfaction with the use of CPAP masks have been frequently reported. CPAP mask interface induced side effects, such as air leak, noise, discomfort, and facial skin problem, considerably affect the overall effectiveness of CPAP treatment [1-6]. Conventional CPAP masks found in marketplace include Full-face mask (Oronasal mask), Nasal mask, Nasal pillow, Oral mask and Total face mask. Table 1 shows the comparison amongst different types of CPAP masks in terms of contact area and air delivery pattern.Table 1.

Table 1 shows the comparison amongst different types of CPAP masks in terms of contact area and air delivery pattern.Table 1. Classification of different types of CPAP maskMask Full-face masks Nasal masks Nasal pillows Oral masks Total face masksContact area Cover both nose and mouth Cover nose only Contact inside rim of the nostrils Cover mouth only Cover entire faceAir delivery path Through nose and mouth Through nose only Through nose only Through mouth only Through nose and mouthMask interface related side effects can be developed, despite of the type of masks used. Basically, CPAP mask selection is decided by clinical effectiveness, patient’s individual characteristics, previous experience of medical devices and economic aspects [1]. Although CPAP masks greatly vary from configurations and materials, mask concept design and size range are mainly based on the averaged individual characteristics. CPAP mask size ranges are461very limited only coming with small, medium.

CPAP mask size ranges are461very limited only coming with small, medium. Mask manufacturers often provide sizing templates to guide patients in mask selection. However, these sizing templates are very simple only considering individual’s 2D facial features. In fact, individual’s 3D facial features, which completely vary between individuals, have significant influence on the mask fitting. Unfortunately, conventional CPAP masks do not consider much on the variation of 3D facial features in mask design. In addition, there is no comprehensive mask selection system suggesting OSA patients to select the suitable mask before CPAP treatment. It is not rare that some OSA patients have more than two CPAP masks with them due to mask unfitting.2.

It is not rare that some OSA patients have more than two CPAP masks with them due to mask unfitting.2. MethodThe combination of Reverse Engineering and Additive Manufacturing offers the possibility to design and fabricate a bespoke CPAP mask which has the ability to fit individual’s physical characteristics.2.1 Digital facial feature acquisitionIn this study, a total number of 20 volunteers were recruited from School of Engineering, Newcastle University after obtaining the ethical approval. Reverse Engineering technique was used in this study to acquire individual’s facial features. A handheld Creaform Go-Scan 3D scanner with an accuracy of 0.1mm was used to acquire the volunteer’s 3D facial datasets, which is a structure-light projection scanner (white light source). During the 3D scanning process, the 3D scanner was revolved around the volunteer’s whole head to obtain the qualified 3D datasets.

During the 3D scanning process, the 3D scanner was revolved around the volunteer’s whole head to obtain the qualified 3D datasets. To well explain the methodology adopted in this work, several volunteer’s facial databases were selected as examples.2.2 Modification of 3D scanned image3D scanning software, Creamform VXelements, was used to edit the 3D scanned facial datasets. The original 3D scanned images may contain several missing datasets and not smooth. Modification of 3D scanned facial datasets was done with Creamform VXelements, remaining the front view of the facial model. A solid 3D digital facial model was then generated after the facial mesh modification.2.3 Measurement of facial feature based on 3D datasetsWith the 3D datasets of volunteer’s facial feature, measurement about dimensions of the key facial features, both of 2D and 3D, can be easily obtained.

These key dimensions measured will provide guidance on further CPAP mask selection and customized mask design.3. Results3.1 Digital facial acquisition and facial mesh modificationIn this study, 20 volunteers were involved, including 15 males, and 5 females. The whole scanning process for each volunteer was approximately about 1 minute. Original 3D scanned image should be edited properly by filling the missing datasets and smoothing the image. Figure 1 shows an example of one volunteer’s 3D facial datasets before and after image editing.People have totally different facial characteristics; some volunteer’s 3D facial datasets may be more difficult to edit compared to others. For example, as shown in Figure 2, this volunteer has462lots of facial hair, which increases the difficult when editing the 3D image. These missing datasets must be fully filled before exporting for further use.Fig 1. An example of 3D facial datasets before (1) and after (2) editingFig 2.

These missing datasets must be fully filled before exporting for further use.Fig 1. An example of 3D facial datasets before (1) and after (2) editingFig 2. An example of 3D facial datasets before (1) and after (2) editing3D facial mesh modification of all volunteers was shown in Figure 3.Fig 3. Facial mesh modification of all volunteers463After facial mesh modification, a solid digital facial model was generated. Figure 4 shows the front and lateral views of the edited 3D digital images of the volunteer’s facial feature.(a)(b)Fig 4. Front (a) and Lateral (b) view of a solid facial model3.2 Measurement of facial featuresMeasurement of facial features can be done easily and precisely with the help of 3D digital facial datasets. In this study, 3D software (3-Matic), was used to measure the dimensions of facial features. Take one of the volunteer’s 3D facial model as an example. The first step is to mark all key facial features on the digital facial model.

Take one of the volunteer’s 3D facial model as an example. The first step is to mark all key facial features on the digital facial model. As shown in Figure 5, a total number of 22 targeting locations were marked covering all key facial features related to mask design and selection. With these marked targeting points, dimensions of facial features, both of 2D and 3D, can be measured.Fig 5. Key facial feature targeting locationsBecause these targeting points were marked manually, therefore, multiple measurements of facial features were done repeatedly considering the deviation contained. For this facial model, measuring processes were conducted 4 times. Dimensions of 2D facial feature indicated in Table 2 were calculated on average.Except for the difference on the 2D facial features, variation of the 3D facial features of individuals presents a higher challenge on CPAP mask design and selection.

Individual’s 3D464facial features are difficult to be measured manually on people’s face. Conventional CPAP mask designed with less consideration of individual’s 3D facial features cannot guarantee a perfect fit for all individuals, despite several sizes of masks can be offered. However, measurement of 3D facial features for individuals can be easily done from the analysis of 3D digital facial datasets. Individual’s 3D facial features, for example, all details of nose structure, can be precisely obtained using 3D measuring software (3-Matics). Figure 6 depicts the 3D digital nose structure of a volunteer.Table 2.

Figure 6 depicts the 3D digital nose structure of a volunteer.Table 2. Measurement of 2D facial featuresItem The distance between eyes (eyelid) The distance between eyes (Centre of eyes) Width of eye The distance between double cheeks Eyebrow to Eye Width of mouth Height of mouth Bottom of the nose to Upper LipDimension (mm) 26.26 72.94 45.26 84.14 27.48 67.94 23.36 16.25This volunteer’s nose was transversely separated with 9 points into three sections to easily collect the detailed characteristics of nose structure, as shown in Figure 6. These three sections are named as top section, middle section, and bottom section. By analysing these three sections, the width of nose and angle of the nose can be collected. The deviation of all dimensions is about ±3mm.Fig 6.

By analysing these three sections, the width of nose and angle of the nose can be collected. The deviation of all dimensions is about ±3mm.Fig 6. The 3D digital dataset of the nose structureTable 3 Measurement of 3D facial featuresItem Length of nose Width of nose Angle of nose Slope of nose (left view) Slope of nose (right view)Dimension (mm) 37.24 Top: 26.26 Middle:26.55 Bottom:33.80 Top:89.33 Middle: 87.17 Bottom:60.51 Top:18.04 Middle: 21.33 Bottom: 33.58 Top:19.34 Middle: 20.17 Bottom: 33.50465Figure 7 shows the details of the slope of nose measured based on the 3D digital datasets. Figure 8 depicts the angles of nose measured from different sections based on the 3D digital datasets.(a)(b)Fig 7. The slope of nose measured from right view (a) and left view (b)Fig 8. The angle of nose measured4. Customization mask design and fabricationFull-face masks have been mostly selected compared to other types of CPAP masks, such as Nasal masks and Nasal pillows.

The angle of nose measured4. Customization mask design and fabricationFull-face masks have been mostly selected compared to other types of CPAP masks, such as Nasal masks and Nasal pillows. Full-face masks are designed to cover both the nose and mouth area, which is ideal for OSA patients who are mouth breather. However, Full-face mask induced side effects, such as air leak, noise, discomfort, and facial skin problems have been reported frequently. If a patient wears an ill-fitted Full-face mask, a gap between facial skin and mask interface can be developed. Unintentional air leak can occur due to the occurrence of this gap, which can affect the effectiveness of CPAP therapy and cause noise. To minimize the air leak and noise, one of the solutions may be tightening the mask strap to minimize the gap. Based on the facial anatomy analysis, nasal bridge has the thinnest facial skin.

To minimize the air leak and noise, one of the solutions may be tightening the mask strap to minimize the gap. Based on the facial anatomy analysis, nasal bridge has the thinnest facial skin. Consequently, facial skin damage can be easily developed around the mask, particularly at the nasal bridge area, when suffering prolonged excessive contact pressure caused by overly tightening the466mask strap. Therefore, conventional Full-face mask design cannot properly fit individuals due to its standard configurations.4.1 Customized Full-face mask concept designThe possibility of generating a customized Full-face mask which is expected to minimize the side effects caused by wearing an ill-fitted conventional Full-face mask was explored in this study.(a)(b)Fig 9.

The profile of designed curve front view (a) and lateral view (b)With the help of 3D digital facial model obtained using reverse engineering technique, customized Full-face mask design was started by sketching a curve onto the digital facial model. The defined curve can be various meeting different requirements. In this study, this specific curve was designed avoiding the facial bony (cheeks) areas, which can minimize the development of facial skin damages when mask strap was being tightened during treatment. By adopting the methodology introduced in this study, there will be no gap between facial skin and mask interface, therefore effectively minimizing the occurrence of air leak and noise. The comfort level with the mask configuration is estimated to be increased. To achieve this curve, 3D CAD software (Inventor) was used. Figure 9 shows the designed curve profile.(a)(b)Fig 10.

The comfort level with the mask configuration is estimated to be increased. To achieve this curve, 3D CAD software (Inventor) was used. Figure 9 shows the designed curve profile.(a)(b)Fig 10. The profile of redesigned curve front view (a) and lateral view (b)The designed curve needs to be redesigned as a symmetry part before creating the entire mask configuration. Figure 10 shows the profile of the redesigned curve.Based on the curve designed, a series of curve sets can be further generated. The aim of sketching these curve sets is to generate the surface of the customized mask configuration. Figure 11 shows the surface generation. The final step is to convert this surface geometry into467a solid part using thicken function. For this case, 4mm was assigned to form this solid part. Figure 12 shows the solid Full-full mask.(a)(b)Fig 11. Surface generation of Full-face mask front view (a) and lateral view (b)(a)(b)Fig 12.

Figure 12 shows the solid Full-full mask.(a)(b)Fig 11. Surface generation of Full-face mask front view (a) and lateral view (b)(a)(b)Fig 12. Development of solid Full-face mask, front view (a), and lateral view (b)4.2 Customized mask concept prototypingConventional CPAP masks are mainly fabricated using injection modelling technique. Using traditional injection modelling technique, mass production can be easily achieved. Conventional masks with standard configuration can be quickly formed. However, these standard products have significant limitations on design configuration, and sizes. In the case of CPAP masks, probably only few sizes can be selected, such as small, medium and large sizes. It is obvious that these standard masks cannot perfectly fit individual’s physical characteristics.(a)(b)(c)Fig13.

It is obvious that these standard masks cannot perfectly fit individual’s physical characteristics.(a)(b)(c)Fig13. 3D printed customized Full-face mask frame front view (a), back view (b), and lateral view (c)468Different from conventional CPAP mask manufacturing technique, additive manufacturing technique builds parts by adding layer upon layer of materials, which provides a great deal of flexibility on product design and fabrication. Using additive manufacturing technique can easily fabricate a complex object which is hard to be generated using traditional manufacturing technique. The higher efficiency presented by using additive manufacturing technique can be seen as another aspect makes it more optimal to fabricate the customized mask frame.Additive manufacturing was used to fabricate these customized mask designs. In this work, Fused deposition Modelling (FDM) printer was used to fabricate this customized CPAP mask frame. The material used to form this part is PLA.

In this work, Fused deposition Modelling (FDM) printer was used to fabricate this customized CPAP mask frame. The material used to form this part is PLA. The total fabrication time was about 2 hours. Figure 13 shows the 3D printed customized Full-face mask frame.5. ConclusionThis paper has explored a possible way in which a customized mask frame can be designed and fabricated fitting individual’s physical characteristics. The combination of Reverse engineering (3D scanning), 3D CAD design and Additive manufacturing offers the opportunity to achieve a customized mask frame. Human factors have been playing the important aspects challenging the mask fitting in CPAP treatment. Conventional CPAP masks cannot properly fit all individuals, because individuals have totally different facial characteristics, both of 2D and 3D facial features.

Conventional CPAP masks cannot properly fit all individuals, because individuals have totally different facial characteristics, both of 2D and 3D facial features. Besides the influence of facial characteristics, in CPAP treatment, symptom related aspects, such as severity of OSA syndrome and previous experience with the use of CPAP mask, should also be considered carefully. In addition, individual preference, for example, sleep and breathing pattern, also challenge the effectiveness of CPAP treatment. Therefore, to improve the performance of CPAP masks, human factors must be considered carefully into the mask design.AcknowledgementsThis project was supported by Newcastle University, School of Engineering PGR grant program 2016.Reference[1] Brill, A.K., 2014. How ventilation. Breathe, 10(3), pp.230-242. [2] Visscher, M.O., White, C.C., Jones, J.M., Cahill, T., Jones, D.C. and Pan, B.S., 2015.

VELOPMENT OF A HYBRID MANUFACTURING CELL;Integration of Additive Manufacturing with CNC Machining K. Boivie¹, K. Sørby², V. Brøtan² and P. Ystgaard¹¹SINTEF Raufoss Manufacturing AS, Trondheim, Norway ²Dept. of Production and Quality Engineering, Norwegian University of Science and Technology, Trondheim, NorwayAbstractThe application of Additive Manufacturing (AM) for production of injection molding tools and tooling inserts enables significant improvements in regards to, for example, product quality and cycle times. However, the AM based production of tools and inserts is most often far from being rapid and the inserts are usually considerably more expensive than conventionally produced. A combination of AM with CNC milling in a hybrid process route allows for the application of each process to the production of the section of the product geometry for which it is most advantageous.

A combination of AM with CNC milling in a hybrid process route allows for the application of each process to the production of the section of the product geometry for which it is most advantageous. However, this approach also multiplies the number of process steps and therefore also the limiting factors and possible causes of failure. This paper describes the development of a Hybrid Manufacturing cell by integration of AM with conventional CNC milling in a robust, streamlined production sequence.IntroductionThe rapid development of additive manufacturing (AM) technology for metallic materials during the first decade of the 21st century has made a number of alternative processes and materials available to the manufacturing industry. However, different processes have different merits, and are therefore perhaps more suitable for different applications.

However, different processes have different merits, and are therefore perhaps more suitable for different applications. Even if there still are a limited number of materials developed for each process, the material properties in parts produced by metallic AM processes are in general very well comparable to their counterparts produced by conventional methods. In combination with the inherent geometric freedom of AM technology, this has made these processes capable of manufacturing products with industrial grade metallic material properties and a freedom of product geometry never preceded in manufacturing technology. However, despite the progress and new possibilities, still the majority of the manufacturing industry hardly regards AM as a serious manufacturing process alternative.

However, despite the progress and new possibilities, still the majority of the manufacturing industry hardly regards AM as a serious manufacturing process alternative. There are many reasons for this; among those is that there is no tradition for designing parts intended for production by AM, neither are there any established standards for process material and quality management. Another, rather fundamental reason is that building parts by AM is not cheap or “rapid” compared to conventional machining, -and the cost increases with building time and the amount of material added to the part. On the other hand, while the cost of conventional machining increases with the amount of material removed, it is considerably faster for shaping massive objects than AM, it has a wider array of alternative materials, it is a very familiar process and is capable of producing parts with a precision and surface finish that is very different from what can be produced by AM.

In the present situation it is one of the most important challenges for the development of AM to gain wider acceptance and usage of this technology throughout all sections of industry, and to do this it is necessary to bridge the apparent limitations of AM and integrate this technology with conventional manufacturing processes.This situation is well known, and in order to bring the benefits of each manufacturing principle together several hybrid manufacturing solutions have been proposed, [1, 2, 3 and 4]. Since one of the most significant limitations of early metallic AM systems was the apparent stair-stepping on the products’ surface, it was an obvious solution to remove this and to generally improve the precision on critical surfaces by finishing machining [1, 5], and whilethe stair-stepping effect is much less significant in today’s technology, this finishing machining is still a desired feature of any hybrid manufacturing solution.

Other solutions have been dividing the CAD model of the intended product in different modules depending on the most suitable manufacturing method, and then assembled the manufactured modules in a subsequent process step [2]. The more recent development in AM technology have allowed some companies to offer a hybrid manufacturing solution, based on a powder bed metal AM system, where the AM part is built directly on top of a CNC (Computer Numerical Controlled) machined part [3]. However this solution is still based on the principle of two different modules which are designed and produced separately, with significant manual work related to operations such as heat treatment as well as part fixation and positioning required between the subsequent parts of the manufacturing process. It is the object of this paper, to present an even more integrated solution to hybrid manufacturing with a streamlined workflow and minimized intermediate operations.

It is the object of this paper, to present an even more integrated solution to hybrid manufacturing with a streamlined workflow and minimized intermediate operations. With the ambition to develop the integrated hybrid manufacturing process into an industrially relevant and practical solution, the development has been based on industrial case studies and practical investigations. The research performed, as well as this paper, is therefore of a very technical nature.Background: Industrial case studiesThe industrial case studies have had the objectives to: a) verify the benefits of AM to relevant manufacturing industries, in this case injection molding industries, b) evaluate present hybrid manufacturing practice, and c) compare hybrid tooling to conventional tooling. The case studies have been carried out in the form of collaboration between a local, Norwegian injection molding companies, a local Norwegian tool making company, and ConceptLaser GmbH, of Lichtenfels, Germany.

ConceptLaser is a producer of metallic AM systems, with a strong background in injection molding and tooling industry, and which therefore has both experience and expertise in designing conformal cooling systems for injection molding tool inserts as well as practical hybrid manufacturing solutions.Case 1: Insert for a bracket to an office chairThe bracket is a critical component that connects the chair lift with the seat and thereby carries a heavy load. To maintain sufficient strength in this connection the previous design had a piece of steel with a machined, conical hole molded in to the bracket. This is, however, a costly solution, and a new design has increased the molded wall thickness in the critical area and replaced the machined steel part with a washer, which would greatly reduce the cost of each bracket. But the increased wall thickness makes this section of the tool a critical point for cooling.

But the increased wall thickness makes this section of the tool a critical point for cooling. The tool consists of one cavity and a core which also has the function of injection nozzle for the tool. The cooling in the larger cavity is of conventional type with drilled, straight holes, but since the core besides being the injection nozzle, and thus a hot-spot in itself, also is the center point of the section where the new design has greatly increased the wall thickness makes the cooling of the core particularly challenging. The original solution to this issue has been core made from Moldmax (a copper alloy tooling material) and drilled cooling channels in a finger pattern. This is the highest level of cooling that can be achieved with traditional tooling techniques, but still not quite satisfactory for the new bracket design. The alternative solution has been a new core built by AM with parallel conformal cooling channels in an umbrella pattern, (see Figure 1A).

The alternative solution has been a new core built by AM with parallel conformal cooling channels in an umbrella pattern, (see Figure 1A). The cooling channels and the core was designed and made in by ConceptLaser and then machined to the final geometry by a local Norwegian tool maker (see Figure 1B - C).ABCFigure 1: Injection molding core with conformal cooling channel, from the left; conformal cooling channel design, core prior to machining, and the finished tool after testing.Results: The cost for machining of the new core was approximately the same as for the original but with the additional cost for AM made the total cost for the new core double that of the original core. Since the production of the new core, in difference to the original, was splitted on two separate locations it is hardly possible to draw any other conclusions regarding building time, than that it has been longer when the core is made by means of AM.

Practical testing of injection molding with the new conformal cooling core resulted in a reduction of the cooling time from 70 seconds (orginal tool, original bracket design with orginal material thicknes) to 48 seconds (conformal cooling core, new bracket design with increased material thickness). Simulations have indicated that the increased material thickness of the product would bring an increase in cooling time by aproximatly 25 seconds using a conventional core. Therefore, by the integration of conformal cooling to a critical core in this tool could the total cycle time be reduced by 20 % instead of having the expected increase by 20 %.Case 2: Manifold for compressed airThis case component is a manifold for compressed air couplings; the design is highly controlled by customer requirements, and therefore does not entirely follow the design rules for injection molding parts.

The component design consist of four slots for push-in air couplings fixed on a base plate for mounting the manifold to an air tank (see Figure 2). This highly unsymmetrical design has made this component troublesome for production, and during injection molding this component experienced significant warpage in the base plate. The former solution to this has been to keep the parts in the mold for 30 sec., remove them while they are still a bit soft, and manually put them in a fixture and then cool in ice water for 90 sec. (see Figure 2).Figure 2: Manifold for compressed air couplings, injection molded in a double cavity tool, and the set-up for warpage correction. This solution has a 2 min cycle time, requires a crew of three operators and the use of a water tank.Research has shown that conformal and directed cooling can reduce both warpage and cycle times [6], and the purpose of this case study was to investigate this possibility.

This is a double cavity tool where each cavity has three sliders and five cores. For the case studywas a new set consisting of one cavity with sliders and cores manufactured. Similar to case 1, this new tool was built in a hybrid manufacturing procedure; the base parts was milled from Orvar Supreme tool steel [7] at a local tool maker company, sent to ConceptLaser for AM of the remaining part with the conformal cooling, and then returned to the local tool maker for machining, grinding and polishing to required precision and surface quality.Results: During test runs the new tool showed a significant improvement in cooling; but the problem with warpage remained. However, if the part was kept in the tool until it had cooled down enough for the plastic to reach a rigid state, the parts came out with satisfactory flat base plate.

However, if the part was kept in the tool until it had cooled down enough for the plastic to reach a rigid state, the parts came out with satisfactory flat base plate. This has made it possible to omit the fixture and water bath, and cut down the crew to one operator, but it has come at the price of an extension in cycle time to 5 minutes. Clearly, despite the advantage brought by integration of conformal cooling this cannot totally compensate for challenges that are built into the design of the product.With hybrid manufacturing the number of process steps and the complexity of the information flow increased. This resulted in a disproportionate relationship between processing lead time and actual production time. However, there were some very practical issues of this approach to hybrid production of tooling inserts observed. For example: with more process steps and involving more different machines multiples the number of set-ups for each part.

For example: with more process steps and involving more different machines multiples the number of set-ups for each part. This in turn enhances the need safety margins and offsets added to the part geometry, which in turn will increase the material that has to be removed in the finishing machining. Conventional milling and AM building are two fundamentally different processes with fundamentally different process requirements that has to be fulfilled before the work piece can enter the subsequent process step.

Even if the production of this mold had been carried out at the same location the set-up time and labor cost required at each process step would have been substantial, and if hybrid manufacturing shall be a more competitive part of an industrial production system there is need for a more integrated process solution.The concept of the hybrid manufacturing cellPrinciple and hardwareThe basic principle of the hybrid cell is to have an integrated cooperation between an AM machine, in this particular case a ConceptLaser M2 Cusing machine (M2), and a CNC milling machine, in this particular case a Deckel-Maho 5-axis milling machine (DM), where each technology is used to produce the parts of the final tooling inserts which it is most advantageous to do so. For a typical injection molding insert this would mean that a base part will be milled from a block of raw material that has good milling properties and is compatible with the AM building conditions.

For a typical injection molding insert this would mean that a base part will be milled from a block of raw material that has good milling properties and is compatible with the AM building conditions. Special attention in the primary milling operation would have to be given to having parallel surfaces on top and bottom and on the surface quality of the upper surface that will serve as a substrate during the AM building. The base part will then be moved to the M2 machine where the section of the insert with the more complex geometry, typically the conformal cooling, is added to the base part.

The base part will then be moved to the M2 machine where the section of the insert with the more complex geometry, typically the conformal cooling, is added to the base part. Finally, after completed AM building, the raw product will be returned to the DM machine for finishing machining to desired dimensions and surface roughness.Workflow for injection molding tool insertsThe principle for the work flow within the hybrid cell as illustrated in Figure 3 present a number of critical intermediate operations that have a strong influence on the practical work; After the raw material has been cut out from stock, it will require a basic pre- machining, producing among other features a bottom surface to specifications for installation on a palette fixation system to hold the work piece during CNC machining and AM building. The pallet is clamped in the DM; a positioning system determines the datum coordinate system before the actual CNC milling.

The pallet is clamped in the DM; a positioning system determines the datum coordinate system before the actual CNC milling. After CNC milling, and before mounting in the M2, the base part and pallet has to be rinsed and free from chips and cutting fluids.Figure 3: Typical workflow for production of injection molding inserts in the hybrid manufacturing cell. Preparatory operations for AM building marked with * would be desirable to keep at a minimum or, if possible, to be avoided altogether.The quality and actual dimensions of the part need to be controlled, and dependent on the material there may be a hardening operation required. The issues of powder handling; filling and leveling the powder prior to the AM building, and rinsing from powder after completed could be a topic for further development of the hybrid cell, but will not be addressed at this stage.

After removal from the M2, the “raw” insert will require a heat treatment to remove tensions and acquire the desired hardness, before it is returned to the DM for finishing machining. As has been pointed out in Figure 3, there are a number of the intermediate operations that would be desirable to keep at a minimum or, if possible, entirely avoid.Base part materialA good rule of thumb is to keep the number of operations as small as possible, and therefore is it the ambition of the development of the hybrid cell to find a working model that keeps the number of operations to a bare minimum while maintaining a practical working order and product quality.

For the practical development of the hybrid cell, this emphasizes the need for an efficient, workable clamping and positioning system, and also raises some critical issues regarding the choice of material and process conditions for the base part.Three different steels have been evaluated as potential base part materials; Orvar Supreme [7] is a traditional tool steel, corresponding to the standards Premium AISI H13 and DIN W.nr. 1.2344 that is very commonly used for injection molding tools on the Nordic market. This steel has also been used successfully for hybrid manufacturing of injection molding inserts described in case 2. Since the steel material used in the AM building corresponds to DIN W.nr. 1.2709 [8], a sample of the same steel has also been included in the evaluation.

Since the steel material used in the AM building corresponds to DIN W.nr. 1.2709 [8], a sample of the same steel has also been included in the evaluation. This is, in difference to Orvar, a maraging steel that is tempered to desired hardness by age hardening at a considerably lower temperature than conventional tool steels: 490°C in comparison to 1020°C – 1030°C for Orvar. This would in practice mean that an intermediate hardening operation between CNC machining and AM building could be eliminated from the process. The third material for this investigation was Marlok C1650 [9], a tool steel that shares the age hardening properties with the AM tool steel and W.nr.1.2709, but is most commonly used for demanding tooling applications such as aluminum die casting tools.For evaluation of the machining properties of these different steels was a sample product geometry that includes a selection of typical features commonly occurring in tooling made from each of the materials, see Figure 4.

Since it would be desirable if the rinsing of chips and cutting fluids could be kept at a minimum, was the machining for this evaluation made without the use of cutting fluids.Figure 5: Sample product geometry for evaluation of the dry machining properties of the potential base part materials.During this test the operator marked his evaluation of the machining properties for producing various features in each steel by giving 1 – 4 points, where 4 points is very good and 1 point is very bad. See Table 1.Table 1 The completed form for evaluation of the dry machining properties for base part tool steel materials.From this test it was found that W.nr. 1.2709 was most difficult to machine and some cutters were also broken while testing. The total score for machining this material was 17 points, which is significantly lower than the alternatives. Orvar and Marlok both scored 24 points, and the machining properties were judged to be similar by the operator.

Orvar and Marlok both scored 24 points, and the machining properties were judged to be similar by the operator. However, Marlok have two principal advantages over Orvar: it has not scored very bad (1 point) for any of the test criteria, and more importantly it shares the hardening conditions with the AM tool steel. Therefore, the hardening operation between the first milling operation and the AM building in the hybrid cell that is required for Orvar, could be eliminated by choosing Marlok for base part material. Marlok is therefore judged to be the most suited tool steel to be used for base part material in the hybrid cell.The quality of the substrate (start surface for the build) during AM building is of critical importance.

The surface roughness should, according to ConceptLaser be no more than Rz4, which equals a Ra value of 0.4 - 0.8 µm, and since the building depends on aenergy transfer by a laser, it is also important that the laser is not reflected from the substrate’s surface. A second sample block was produced to investigate the possibility to mill a surface that could fulfill these requirements, see Figure 5. Ten different grooves were milled with different cutters, feed and spindle speed. No cutting fluid was used for this test either.Figure 5: Sample block for investigation of milling surfaces.Figure 6: Mahr Perthometer M2 and the milling surface testing sample block.As can be seen in Figure 5 there was not a dramatic difference between the surfaces produced with the different milling parameters. Surface roughness measurements with a Mahr Perthometer M2, see Figure 6, gave a roughness span of Ra = 0.3 - 0.6 µm which fulfils the requirement.

Surface roughness measurements with a Mahr Perthometer M2, see Figure 6, gave a roughness span of Ra = 0.3 - 0.6 µm which fulfils the requirement. However, as also can be seen in Figure 5, all surfaces were rather reflective to light, which would not be optimal in respect to laser absorption. There are still some differences and the most promising groove was milled with a 10 mm end mill and a cutting speed of 50 m/min, a feed of 0,111 mm/tooth and a spindle speed of 1500 rpm. The outcome will change with different parameters and tools, therefore these figures should not be considered as absolute, and it is not impossible that a less reflective surface could be achieved with further experiments.

Otherwise, the alternatives are to either glass/sandblast the surface or to apply a fine coating that will reduce the reflection and evaporate when heated by the laser; one possibility could be to paint the substrate with a black color marker.Clamping and fixation: Positioning and precisionThe M2 machine is designed with two separate stations; one for the manual powder handling and preparation for setting up a build, such as installation of the substrate (base plate or base part), and one process station for the actual building process. The build- and material, containers are set in a carrier system that transports these between the powder handling area and process station.

The build- and material, containers are set in a carrier system that transports these between the powder handling area and process station. Since the positioning of the carriage in the process station must be accurate in order for the laser zero point to remain the same as the set-up zero point in the build chamber and to maintain consistency from one build to the other, it is important to determine the deviations that could be expected due the movements of the carrier. For this purpose was the deviation of the build plate’s position measured during a series of repeated carrier movements. The tests were set up with a solid block of metal fastened to the build plate, and a digital indicator gauge with micron accuracy located in the process station took measurements of the position of the metal block. This test series involved 33 measurements in each axis, where the carrier was moved all the way to the handling station before returning to the process station for a new measurement.

This test series involved 33 measurements in each axis, where the carrier was moved all the way to the handling station before returning to the process station for a new measurement. A second test series with measurements in the X- and Y- directions was also performed. For these measurements was only the base plate moved 140 mm up and down. These tests produced smaller deviations than the tests involving the full carrier movement. This seems reasonable, and indicates that what was being measured in the full movement tests was the total accuracy of the carrier positioning. The Z-axis zero- position of the build plate movement was reconfigured in the handling station before every new measurement was taken, see Figure 7. The results of these tests show that on most occasions did the carrier positioning exceeded the gauge accuracy.

The results of these tests show that on most occasions did the carrier positioning exceeded the gauge accuracy. Based on these results it is a reasonable conclusion that the positioning of the carrier has an insignificant effect on the overall machine accuracy.Figure 7: Deviations in the carrier positioning.AM build accuracy on the substrateAfter a regular service on the M2-machine it was found that there were still some geometrical deviations in produced parts. These deviations would lead to the need to apply additional surface offsets for the post-machining, which would make the required post machining grow both in regards of the surface area that needs to be machined, and how thick the material of the finishing machining area would be [4]. As has been concluded previously in this work and also been observed in other research [10], is the machining properties of the AM steel material (identical to W.nr. 1.2709) not attractive.

As has been concluded previously in this work and also been observed in other research [10], is the machining properties of the AM steel material (identical to W.nr. 1.2709) not attractive. It is therefore highly desirable to keep the required finishing machining as little as possible, in particular since in the hybrid manufacturing process this would have to be done when the steel is in hardened condition. Therefore, a method for finding and improving the build accuracy has been developed.The starting point for this method is a grid of 10mm high cylinders that is built directly on the substrate. The cylinders are measured by means of a high precision coordinate measuring machine (CMM), and the approximate center point location is found for each cylinder by the use of the maximum and minimum value in the x- and y-axis of the cylinder.

The next step is to make an area for each center point by determining the limits where the actual center point lays plus-minus one third of the known radius. The measurements are then subjected to a sequential quadratic programming optimization [11, 12, and 13], where the smallest radius in the grid is set to be as big as possible. In this optimization non-linear equality constraints are used to establish the known grid for the supposed orientation of the cylinders. By using the grid points found it is possible to calculate the actual desired center point of each cylinder and then determine the offsets from this point in each axis. By plotting these offsets with their respective positions, and curve fitting the deviations, it is possible to find the equations that correct the laser’s positioning.

By plotting these offsets with their respective positions, and curve fitting the deviations, it is possible to find the equations that correct the laser’s positioning. Implementation of this test in the present M2 machine has reduced the maximum negative deviation to a sixth of the previous recorded value.Clamping and fixation: CNC millingSince there is an Erowa clamping system [15] pre-installed in the M2 machine from delivery this has been the primary system for investigation as the generic clamping and positioning system in the hybrid cell. The Erowa system (see Figure 8) consists of “PowerChuck P” clamping unit which is to be mounted in the machine, and movable pallets that are held firmly by the clamping units during processing of the work piece which will be fixed to the pallet. The pallet is clamped to the chuck with a clamping force of 10 000 N, generated by springs which can be released by compressed air, (see Figure 8).

The pallet is clamped to the chuck with a clamping force of 10 000 N, generated by springs which can be released by compressed air, (see Figure 8). According to the manufacturers specification the repeatability of the position of the pallet to the PowerChuck is 0.002 mm. This accuracy of positioning is within the expected offset values for AM building and should therefore be sufficient for the subtractive and additive processes in the hybrid cell. However, the stability of the work piece holding system could be a limiting factor for the efficiency of the subtractive operations in the cell and therefore it is a critical issue whether the Erowa clamping system is able to withstand the forces and maintain the stability and position accuracy during the milling operation.Figure 8: Erowa PowerChuck P [15] 158 x 230, for mounting onto the work table in the milling machine, and a movable pallet, Ø148 for work piece fixation. The PowerChuck release function is operated by compressed air.

The PowerChuck release function is operated by compressed air. Automatic air supply can be made available in the DM machine.A series of milling tests has been performed in order to investigate the stability of the clamping system. For these tests has the work piece material been a mild steel with a hardness of 260 HB, (approximately the same as HRc 27), which is comparable to unhardened tool steel; -Marlok is delivered in soft annealed condition at HRc 28 – 32. The milling tool was a Ø20 end mill (ISCAR HM90 E90A-D20-3-C20) with indexable carbide inserts. The tests were carried out as side milling and slot milling in a Bridgeport VMC milling machine.

The milling tool was a Ø20 end mill (ISCAR HM90 E90A-D20-3-C20) with indexable carbide inserts. The tests were carried out as side milling and slot milling in a Bridgeport VMC milling machine. The properties of this machine (stiffness, spindle design, etc.) are comparable to the properties of the DM machine that will be used in the hybrid cell and the working space of the DM is too small to carry out the necessary measurements, see Figure 9.Figure 9: The Erowa PowerChuck and pallet set up for machining test in the Bridgeport VMCFigure 10: Cutting fluid penetrates the pallet and chuck.The cutting data used in the tests are shown in Table 2.

The material removal rate under these conditions is satisfactory for a rough machining operation on work pieces of the size that will be processed in the hybrid cell.Table 2 Test data for the milling testsDepth-of-cut, ap Engagement, ae Spindle speed, N Cutting speed, vc Feed speed, vf Feed per tooth, fz Material removal rate, MRRSide milling 5.0 mm 4.0 mm 6000 r/min 380 m/min 1800 mm/min 0.1 mm 36 cm3/minSlot milling 3.0 mm 20.0 mm 3000 r/min 190 m/min 900 mm/min 0.1 mm 54 cm3/minDue to reduced stiffness towards the ends of the work piece was machining on the work piece’s unsupported ends expected to be more challenging than machining close to the center of the clamping. However, with the chosen cutting parameters, there was no apparent difference in stability or vibrations between the center and the end of the work piece.

However, with the chosen cutting parameters, there was no apparent difference in stability or vibrations between the center and the end of the work piece. It seems that it is the machine and cutting tools in themselves rather than the clamping system that would be the limiting factors for the material removal rate. No large vibrations have been observed during any of these tests. It can therefore be concluded that the Erowa clamping system most likely is sufficiently strong and stable for the machining tasks expected in the hybrid cell. In order to ensure that these tests should be limited by the pallet systems stability rather than by the increase in temperature during slot milling were these tests carried out using cutting fluid. It was observed that cutting fluid penetrates the pallet as well as the chuck and that some remains at the underside of the pallet, see Figure 10.

It was observed that cutting fluid penetrates the pallet as well as the chuck and that some remains at the underside of the pallet, see Figure 10. This is not critical for the function of the chuck or pallet; however, it stresses the point that if cutting fluid is used, then both the pallet and base part must be thoroughly cleaned before the set-up transferred to the additive machine in the hybrid cell.The Erowa clamping system performed well in the milling tests. However, if it should occur that the stability of the machine tool and/or clamping system would become a limiting factor, there are actions that could be taken to improve the milling operation without changing the present set-up:High speed machining: In high-speed machining, high cutting speed is used, while the size of each chip is relatively small, and therefore the cutting force is also small. The DM milling machine that will be used in the hybrid cell has a maximum spindle speed of 18 000 rpm.

The DM milling machine that will be used in the hybrid cell has a maximum spindle speed of 18 000 rpm. By exploiting the high spindle speed, a relatively high material removal rate could be achieved, without increasing the cutting force.Rough milling before mounting the work piece to the pallet: In cases where the clamping system is too weak to permit satisfactory material removal rate, it is a possible solution to carry out the roughing operation in a separate setup before the work piece is mounted to the pallet. This could also be a solution to particularly challenging machining where cutting fluids would be necessary. It would introduce an additional procedure before the work piece is mounted to the pallet and processed in the actual hybrid cell. However this is likely to be comparably small extra work besides what is necessary anyhow.

However this is likely to be comparably small extra work besides what is necessary anyhow. The work piece will have to be cut out to size from stock material, the bottom surface will have to be milled to specified flatness before it can be fastened to the pallet, and there is need for some feature that can be used for the fixation of the work piece to the pallet. If any of these operations require the use of cutting fluids, which the cutting of the work piece from stock do, or if the work piece is not sufficiently clean to begin with, which is unlikely, the work piece will have to be cleaned at least at one occasion anyhow. It is more work saving to do this before the processing in the hybrid cell has started, than to break in this process and clean both work piece and pallet.

It is more work saving to do this before the processing in the hybrid cell has started, than to break in this process and clean both work piece and pallet. Since this has to be done, it makes perfect sense to have done the operations that are most critical for the need of cleaning beforehand and thus possible save one cleaning operation during the hybrid cell processing.Conclusive remarks and further workThe concept as well as the working principle of the hybrid manufacturing cell are under development and will continuously be further refined. Still, the present investigations and evaluations have confirmed the feasibility of the project and indicated that there are significant advantages to be gained from this system as compared the common solution to hybrid manufacturing that is used today. Evaluation of the machinability and milling conditions has concluded that Marlok C1650 is a tool steel with attractive properties for hybrid manufacturing of injection molding tool inserts.

Evaluation of the machinability and milling conditions has concluded that Marlok C1650 is a tool steel with attractive properties for hybrid manufacturing of injection molding tool inserts. However, since the use of cutting fluids is undesirable in the hybrid cell, while some features of a tool insert may be very challenging to mill without cutting fluid it may be needed to perform a preparatory machining with cutting fluid, followed by cleaning, of the work piece before it is attached to the palletand processing in the actual hybrid cell can start. Flat milling of the tool steel that would be used as base part material in the hybrid cell can be set to fulfill the requirements for surface roughness; however it is likely to produce a surface that may require some extra treatment to ensure satisfactory laser absorption. The precision of the Erowa pallet system and the movements in the ConceptLaser M2 Cusing machine are not limiting factors for the precision of the hybrid cell.

The precision of the Erowa pallet system and the movements in the ConceptLaser M2 Cusing machine are not limiting factors for the precision of the hybrid cell. An improved system for determining and compensation of the laser deviation has been developed, which has resulted in greatly improved precision in the ConeptLaser M2 Cusing machine that will be used in the hybrid cell. The Erowa PowerChuck and pallet system is rigid and strong enough to withstand the loads expected during the milling operation. However, should a situation occur when the stability of this system would be unsatisfactory, there are actions that could be made to compensate for this with the present system and setup.

An Erowa PowerChuck and palette system has been installed in the Deckel- Maho 5 axis milling machine, which together with the measuring probe and the preinstalled Erowa chuck in the ConceptLaser M2 Cusing machine, provides basic physical conditions for integrated operation of the hybrid cell.Acknowledgements The development of the hybrid manufacturing cell is being supported by a collaboration of several national and international research projects and research programs; 7th FP EU project IC2 ,”Intelligent and Customized Tooling”(NMP-2009-4.0-5, Grant agreement no: 246172), NFR (Research Council of Norway) project HYPRO, and CRI Norman (Center for Research driven Innovation), have all contributed to the progress so far.References[1] integration of additive and subtractive processes for hybrid layered manufacturing”, Robotics and Computer-Integrated Manufacturing, Volume 26, Issue 5, October 2010, Pages 490-499, ISSN 0736-5845, DOI: 10.1016/j.rcim.2010.03.008.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDevelopment of a Hybrid Manufacturing process for Precision Metal PartsMissouri University of Science and TechnologyLaser Aided Manufacturing Processes LaboratoryDepartment of Mechanical and Aerospace EngineeringLeon Hill, Todd Sparks, and Frank Liou1 AbstractThis paper summarizes the research and development of a hybrid manufacturing process to produce fully dense metal parts with CNC-level precision. High performance metals, such as titanium alloys, nickel superalloys, tool steels, stainless steels, etc. can beneﬁt from this process. Coupling the additive and the subtractive processes into a multi-axis workstation, the hybrid process, can produce and repair metal parts with accuracy. The surface quality of the ﬁnal product is similar to the industrial milling capability.

The surface quality of the ﬁnal product is similar to the industrial milling capability. To achieve such a system, issues of the metal deposition process and the automated process planning of the hybrid manufacturing process will be discussed.2Introduction and BackgroundIn this paper, a hybrid manufacturing system using a laser metal deposition (LMD) process is integrated with subtractive CNC milling creating a hybrid manufacturing cell. Figure 1 is a high level functional decomposition of a hybrid system. The system description can be decomposed into 4 discrete phases, as described below:1935User Information InputMotionSubtractive PlanningPartMaterialsSubtractive ProcessAdditive PlanningAnalysisUtilitiesAdditive Process1234Figure 1: Functional Decomposition of Hybrid Manufacturing System1.

Inputs The ﬁrst phase denotes the inputs into the process.User Information Input User information is what describes part geometry, a CAD model, and knowledge taken from pervious experience. These are the independent variables that the user has most control over.Materials The materials selected for the process will determine certain parameters. The user can select substrate and powder material. If the user is performing a repair, then the orignial part’s material becomes a constraint.Utilities Hybrid systems require inputs that do not directly contribute to the outputpart, such as electricity, shielding gas, or machining coolant.2. Process planning The second phase denotes the additive and subtractive planning is performed by CAM software and simulation software. The plan must include any premachined areas, typically necessary with part repair, and post machining after deposition [1].

The plan must include any premachined areas, typically necessary with part repair, and post machining after deposition [1]. Subtractive Planning The subtractive planning step is conventional CNC machineplanning.Additive Planning For additive planning, inﬁll patterns, laser approach directions,powder feed rate, and part geometry will deﬁne the path plan.3. Manufacturing Phase 3 is the manufacturing processes. Additive processes typicaly have poor surface ﬁnish. It is key for a hybrid manfuacturing system (HMS) to ensure the two processes do not interfere. The ability to integrate the additive and subtractive portions across all processes is key to a successful rapid high precision hybrid system [2]. Motion There can be various tactics for moving the workpiece and/or the tool.

Motion There can be various tactics for moving the workpiece and/or the tool. Typ-ical motion systems used include both CNC machines and robot arms.Subtractive Process The subtractive process is typically considered to be milling, but other processes, such as grinding, are a distinct possibility. Machining may be performed before and after the deposition process and is traditional CNC G- code [3].1936Additive Process For the purposes of this discussion, blown powder laser deposition of metals is discussed. However, many other additive processes can be viable candidates for the hybrid treatment. When integratign with an existing motion system such as a CNC milling machine, custom M-codes must be added to denote when to turn the laser on and oﬀ, the power density, and powder delievery [3].4.

Outputs The ﬁnial phase is the output products of the hybrid system.Part This is the desired output of the manufacturing activity.Analysis If the part is unsatisfactory, the user inputs must be modiﬁed. Even if the part is within desired tolerances, information from the part production or repair process can be a useful feedback.3 Tool Path Planning3D printed parts, metal or plastic, are printed by ﬁrst laying down an outline of each layer, then ﬁlling in the layer. The object must be sliced by software prior to a ﬁll pattern path plan. Building overhangs and other complex geometries sometimes require building additional support structures. Support structures would have to be machined away later adding time and cost to the process. Using multiaxis machine capabilities, the build plan can rotate so the deposition direction is always vertical.

Using multiaxis machine capabilities, the build plan can rotate so the deposition direction is always vertical. Figure 2 illustrates how multi-axis geometry can be utilized to avoid creating support structures [3].Figure 2: (a) Build part with support structure; (b) Rotation for for next portion of build (c) Post rotation, build component in diﬀerent direction [3]Path planning must also take into account uniform layer thickness and post machining path. Some circumstances may require intermediate machining processes if geometry will not allow for machining after a certain point in the LMD process. By integrating the subtractive and additive processes, machinning steps can be performed trivially saving time, money and material costs.1937Figure 3: Multi-Axis Adaptive Slicing [4]Figure 3 shows slices that do not have a uniform thickness. As the part must curve as it proceeds, one side of the layer must be thicker than another.

As the part must curve as it proceeds, one side of the layer must be thicker than another. Since non-uniform layers are very diﬃcult to control in additive, a uniform section comprised of several layer is deposited. Then an intermediate machining process cuts the uniform section into a non uniform section creating the desired contour. This methodology also aids in prevent tooling crash during machining [4].Coverage toolpath planning is very critical to depositing fully-dense metallic parts. The techniques applied to each layer eﬀect the ﬁnal product. A poorly chosen ﬁll pattern will result in voids or over building in the ﬁnal part. In order to account for variations in part geometries, an adaptive ﬁll pattern is best suited to this task.Figure 4: Diﬀerent types of commonly seen voids in deposition [5]Part geometry and poor tool path planning will lead to voids. Figure 4 illustrates common voids that occur within a part.

Figure 4 illustrates common voids that occur within a part. Void type I occurs when the area to be ﬁlled is too large for a single pass, but not large enough for another parallel oﬀset path. Void type II occurs when corner geometry has too sharp of an angle. The angle is considered too sharp depending on speciﬁc process parameters such as laser spot size. Void type III will occur if the oﬀsetting algorithm generates more than one loop allowing a void between the separated loops.Proper pattern planning can prevent voids from being deposited into the part. Figures 5 and 6 show the two common types of ﬁll patterns. In ﬁgure 5 (a) is a conventional raster pattern1938while (b) alternates direction each layer. Figure 6 illustrates two types of parallel oﬀset ﬁll patterns. The red circles denote locations where possible voids may occur. Each pattern has a geometry where it will best utilized. However, the goal of hybrid manufacturing is automating as much of the process as possible.

Each pattern has a geometry where it will best utilized. However, the goal of hybrid manufacturing is automating as much of the process as possible. Therefore, software must be able to determine the best pattern for a given geometry. Furthermore, the ﬁll pattern should adapt to variations in the part geometry.(a) Fixed direction(b) Interlaced directionsFigure 5: Diﬀerent raster toolpath generation patterns [6](a) Contour-parallel oﬀset pattern(b) Spiral oﬀset patternFigure 6: Deposition void occurrence for two patterns. [6]In order to automate the planning process, possible void locations need to be addressed by software. An algorithm can be used that reliably predicts where voids will occur. The process works by comparing the area to be ﬁlled on two consecutive passes and takes the diﬀerence. The diﬀerence is compared to a user deﬁned tolerance. If the diﬀerence is greater than the allowed tolerance, the process planning software will use a diﬀerent ﬁll pattern.

The diﬀerence is compared to a user deﬁned tolerance. If the diﬀerence is greater than the allowed tolerance, the process planning software will use a diﬀerent ﬁll pattern. Unfortunately, the tolerance required varies by material and process parameters. Thus the tolerance must be empirically determined on each machine for each material. Figure 7 illustrates the adaptive process [6].1939Figure 7: Adaptive toolpath for complicated geometry [6]4 Additive Dimensional FeedbackAn additive process must produce shapes as close to the CAD model as possible. The more accurate the process is, the less post machining is required allowing further savings on time and resources. This section quantiﬁes the dimensional accuracy obtainable from LMD with closed loop height feedback control.A part made from an additive process is built in layers. The layer thickness ﬂuctuates throughout the part build.

The layer thickness ﬂuctuates throughout the part build. The layer thickness depends on the melt pool, scan speed, material mass ﬂow rate, and laser power. The melt pool is a region where the material is molten. Variations in layer thickness are diﬃcult to predict in simulation [7]. Layer thickness is thin compared to overall build dimensions. Therefore, most require hundreds of passes creating large error propagation.Implementation of a sensor allows for closed loop feedback on build height. Using a pyrometer the machine can track the position of the melt pool. The measured value is compared to a reference value. The controller determines a boolean value for the melt pool position. If the melt pool is in the wrong location the machine feed rate will slow down. This allows the more material to build.

The controller determines a boolean value for the melt pool position. If the melt pool is in the wrong location the machine feed rate will slow down. This allows the more material to build. Once the controller determines the melt pool to be in the proper location it will speed back up.Planck’s law, seen in wavelength form in equation 1, describes the incadsecent radiation from an blackbody at a particular wavelength and temperature. Figure 8 illustrates Planck’s equa- tion at selected wavelengths for a range oftemperatures are plotted. The pyrometer works at .45 µm and 1.8 µm. Since the laser used for deposition is 1080nm, there is no interference.19402100 K900 K2500 K1.0Spectral Intensity (Wm2sr)0.220001500 K0.61300 K2300 K10000Wavelength (nm)1100 KPhoton Emissions from the LMD Process4000500 K80001900 K0.8700 K00.41700 K60000.01e17As seen in 8 the emittance drastically increases with temperature.

Getting a precise tem- perature measurement with an optical pyrometer in an area of such high thermal gradient can be diﬃcult, but for the purposes of this work, the sensor only needed to determine the presence or absence of the melt pool.Figure 8: Spectrum radiance wavelength for temperatures2hc2 λ51 λkB T − 1Bλ(λ, T ) =(1)hcewhere Bλ is the spectral radiance. λ is the wavelength. T is absolute temperature. c is the speed of light. h is Planck’s constant.The performance of a deposition system with and without a sensor can be trivially demon- strated. Figure 9 shows three parts built using the aforementioned pyrometer. Figure 10 is a part made using the same parameters as ﬁgure 9 but the layer thickness was not as simulated. Thus after several passes the melt pool was no longer on the part build surface. This demonstrates that height control can correct for improperly set parameters.

Thus after several passes the melt pool was no longer on the part build surface. This demonstrates that height control can correct for improperly set parameters. The sensor does not ensure maximum eﬃciency, but can account for unknown materials and slight errors from simulation to practice.1941Figure 9: From left to right, sample parts made with with height feedback control at 20mm, 15mm, and 10mm [8]Figure 10: Parts made without height control system. The standoﬀ distance increased quickly after a few layers. The part stopped being built up after the powder stream com- pletely missed the melt pool. [8]The parts built with height feedback control were built to heights of 10mm, 15mm, and 20mm. All parts had similar lengths and widths. Length diﬀerences were within 0.53 mm, and width diﬀerences were within 0.18 mm. The height diﬀerences were 0.22 mm in 10 mm parts, 0.94 mm in 15 mm parts, and 0.23 mm in 20 mm parts.

Length diﬀerences were within 0.53 mm, and width diﬀerences were within 0.18 mm. The height diﬀerences were 0.22 mm in 10 mm parts, 0.94 mm in 15 mm parts, and 0.23 mm in 20 mm parts. The diﬀerences between the1942set height and the actual height were +- 0.12mm in 10 mm parts, +- 0.49 in 15mm parts, and +- 1.02 mm in 20 mm parts [8].

The diﬀerences between the1942set height and the actual height were +- 0.12mm in 10 mm parts, +- 0.49 in 15mm parts, and +- 1.02 mm in 20 mm parts [8]. In table 1 the build data is summarized.Table 1: Accuracy and Repeatability Test Results [8]Run #1 2 3 4 5 6 7 8 9Part Heightmm 10 10 10 15 15 15 20 20 20Measured Heightmm 9.9 9.88 10.1 14.51 14.92 15.45 21.02 20.93 20.79Measured Lengthmm 22.05 21.85 21.78 21.78 21.25 21.79 22.02 21.84 21.72Measured Widthmm 1.94 1.96 1.9 1.92 1.96 2.05 1.93 1.99 1.87Build Times 391.81 392.95 401.82 593.42 610.03 644.01 821.93 916.94 922.44Diﬀerence from the set height mm -0.1 -0.12 0.1 -0.049 -0.08 0.45 1.02 .93 .79Raµm 93.05 106.35 116.85 98.9 99.55 99.15 100.2 98.85 97.05 Machining and Veriﬁcation of Deposited PartIn order to fully automate the hybrid process, hardware must be implemented to close the loop. The part dimensions must be within tolerance.

The part dimensions must be within tolerance. Post machining is used to ensure a ﬁnished part has both the desired surface ﬁnish and dimensions. An acoustic sensor can be used to provide depth-of-cut (DOC) feedback during the post machining process. The sensor helps ensure the ﬁnal part is within tolerance and provides feedback on the deposition.Figure 11: Experimental Components for Acoustic Emissions Sensor [9]The acoustic sensor implements hardware and software to accurately measure the DOC during post machining. Figure 11 illustrates the sensor components. A neural network is1943trained to take data from the acoustic sensor and process it. The camera is used to observe the status of the cutting tool. Its sharpness aﬀects the acoustic emissions. This allows the neural network to adjust for variations in tool sharpness.

The camera is used to observe the status of the cutting tool. Its sharpness aﬀects the acoustic emissions. This allows the neural network to adjust for variations in tool sharpness. It cannot adjust dynamically as the camera can only analyze the cutter before and after cuts.In order to ensure accuracy of the sensor calibration cuts with known depths of cut were taken. After machining the specimen was 3D scanned and analyzed. Figures 12 and 13 show the 3D scan and the correlation between DOC and measured DOC respectively [9].Figure 12: Scanned Deposited Material (a) Machined (b) Original (c) Removed [9]This provides proof of concept for the acoustic emissions sensor. It allows for dynamic feed- back for DOC during post machining. The sensor can be utilized to ﬁnd portions of the deposit that were under or over deposited. This method can be used to provide the user with information about their deposit. Furthermore, this sensor can communicate with the controller.

This method can be used to provide the user with information about their deposit. Furthermore, this sensor can communicate with the controller. If a portion is under deposited, the sensor can communicate where another depo- sition pass should occur. Implementation of the acoustic sensor allows for more automation and less human interaction.1944Figure 13: Measured Depth-of-cut from the Sections and Detected Depth-of-cut for a De- posited Material Detected by the Acoustic Emission Sensor [9]6 Part Repair Using Hybrid ManufacturingPart repair technology is a large area of focus in both military and industrial applications. As shown in ﬁgure 14 there are four main categories of damage: pitted, cracked, worn out surface, and broken.(a) Pitting(b) Cracking(c) WearFigure 14: Damages types [10]Sizes of the damage is used to classify the type of damage. Pitted is considered large length and width compared to depth.

Pitted is considered large length and width compared to depth. Cracked is considered very small width but relatively long length. Worn out surfaces are very shallow and occur at points of contact between moving or vibrating parts.1945Figure 15: Repair Process [11]Repairing a part will entail machining away the damaged area, depositing over the machined portion, post machining, then if necessary a specialized ﬁnish procedure such as grinding and polishing. The paths may be determined by Minkowski Sum and Subtraction [10]. The part must be carefully aligned to match CAD models. A touch probe with in house software can align the part allowing comparison and accurate repair [11]. On going research uses a 3D scanner and software to generate a Voxel model. The Voxel model is overlaid with a CAD model which is analyzed by software to determine the damaged regions.For controlled repair experiments, samples were prepared with minimal machining and/or a deformation force of 3500N.

The defects were then repaired with LMD hybrid process. The repair process requires the defected area is clear of any foreign particles. The process for LMD had to be precisely planned using software. Heat transfer calculations were needed to determine time for deposition and preheating. Simulations also help ensure minimal heat aﬀected zone (HAZ). The repaired regions were cut out into tensile test specimens and tensile tested. The results are summarized below.1946(a) Comparative Results of UTS(b) Comparative Results of YS(c) Comparative Results of % ElongationFigure 16: Mechanical testing of laser deposited Ti64 [12].Results show the repaired Ti64 specimens conform to ideal properties of Ti64. The largest changes in properties was percent elongation. Titanium’s ductility is largely dependant on the microstructure and grain size. Thus the deposited microstructure will diﬀer from the ideal substrate microstructure.

Titanium’s ductility is largely dependant on the microstructure and grain size. Thus the deposited microstructure will diﬀer from the ideal substrate microstructure. The tensile test results show a very close relationship in UTS and YS demonstrating the mechanical properties of repaired samples have not been altered to a large extent [12].7 Mechanical Properties and Microstructure of HybridManufactured PartsHigh incident energy, high laser power low travel speed, result in slow cooling rates. Faster cooling rates occur when the laser power is low and the travel speed is higher. For LMD, rapid solidiﬁcation with an ultra-high temperature gradient is common. With properly selected deposition parameters, an ultraﬁne microstructure is possible, which results in a more uniform distribution of the components. The cooling rate is directly related to the size of the substrate, layer thickness, and total number of layers deposited.

The cooling rate is directly related to the size of the substrate, layer thickness, and total number of layers deposited. New layers will melt and reform the microstructure of previous layers up to 14mm below the melt pool.In ﬁgures 17 (a) and (b) illustrate the diﬀerence between the microstructure within the same deposit. The lower layers experience a slower cooling rate compared to the higher layers.1947Therefore, the slowing cooling allows for more coarse dendrite grains to form while the faster cooling rate creates ﬁner dendrite grains [13](a) Top layers(b) Bottom layersFigure 17: 316 Stainless Steel microstructure, diﬀerent positions within the laser deposition zone. [13]Cooling rate has large impact upon the mechanical properties and microstructure. For Ti- 6Al-4V, as the cooling rate decreases the α lath thickness is known to increase. The β transus plays a crucial role in determining the microstructure. The Vanadium provides the β stabilization.

The β transus plays a crucial role in determining the microstructure. The Vanadium provides the β stabilization. Cooling of the β phase determines the grain boundaries. Air cooling results in ﬁner needle-like α phase while slow cooling forms coarse plate like α phase. Figure 18 illustrates the diﬀerence in microstructure on cooling rate.1948Figure 18: SEM micrographs of the Ti-6Al-4V deposit processed by laser deposition, (a,b) quicker cooling rate (c, d) slower cooling rate [14]The mechanical properties diﬀer depending on the properties and phase amounts present in the microstructure. With increasing the cooling rate, the width of the primary α lath de- creases the aspect ratio and volume fraction of primary α increases. This increases hardness and tensile strength while decreasing the ductility.

This increases hardness and tensile strength while decreasing the ductility. Table 2 shows the results of the tensile test according to cooling rate [14].Table 2: Summarizing Tensile Test Data by Cooling Rate [14]Cooling Rate1500 C/S5500 C/SUTS (MPa) YS (MPa) Hardness (HV) UTS(MPa) YS (MPa) Hardness (HV)Sample 1 Sample 2 Sample 3845 890 850825 835 830415 425 4201120 1190 1110995 1140 985405 410 400Typical grain size of laser deposited Ti-6Al-4V ranges between 100µm to 600µm. However, a friction stir process (FSP) can produce a forged like microstructure. The FSP modiﬁes the base material microstructure for a highly reﬁned α grain microstructure. A band 100µm wide did not undergo a β transformation was also observed. This 100µm band was ob- served around the dilution zone formed from laser interaction.

A band 100µm wide did not undergo a β transformation was also observed. This 100µm band was ob- served around the dilution zone formed from laser interaction. The FSP reﬁned Ti-6Al-4V is expected to increase fatigue life by delaying the fatigue crack initiation [15].19498 ConclusionsA summary of the development of a hybrid manufacturing process incorporating both laser metal deposition and CNC machining was presented. Much of the eﬀort presented was fo- cused on the repair application, as the hybrid process presented here has a distinct advantage over other metal additive processes when applied to repair:The blown powder laser deposition process is not constrained by a powder bed, so it can begin processing with arbitrary pre-existing geometry.The CNC machining process can produce a ﬁnal surface ﬁnish not constrained by powder particle size, layer thickness, or deposited track width.References[1] K. Lorenz, J. Jones, D. Wimpenny, and M.

Amine, J. Newkirk, and F. Liou, “Methodology for Studying Eﬀect of Cooling Rate During Laser Deposition on Microstructure,” Journal of Materials Engineering and Performance, 2014.[15] R. Francis, F. Liou, and J. Newkirk, “Investigation of Forged-Like Microstructure Pro- duced by a Hybrid Manufacturing Process,” Solid Freeform Fabrication Symposium.1951.

velopment of a low cost imaging system for a laser metal deposition processShyam Barua1, T odd Sparks2, F rank Liou3 Department of Manufacturing Engineering Missouri University of Science and Technology barua1223@gmail.com1, toddesparks@gmail.com2, liou@mst.edu31 AbstractThe size of the melt pool created by the laser is one of the most important quality characteristic in a laser metal deposition process. This paper discusses the development of a low-cost vision system to automatically determine the size of the melt pool for in-process control.To cope with the intense infrared signal from the laser and melt pool, external ultraviolet illumination is paired with narrow bandpass ﬁlters on a usb microscope to achieve a clear image of the melt pool.

The sensitivity of the melt pool to changes in system parameters and various substrate materials are also evaluated.2IntroductionLaser metal deposition is a type of additive manufacturing process in which a laser is used to create a melt pool to which additional material (typically metals) is added to build up a part. The material to be added can be transferred onto the substrate by powder injection,preplaced powder or by wirefeeding [1]. This process,in conjunction, with a computer numerical control (CNC) motion system ,enables us to build three dimensional parts which are functional in their intended purpose.The various main process parameters are incident laser beam diame- ter,process speed, material feed rate.The other factors involve gas ﬂow rate,depth of focus and laser absorbtivity of substrate[2].The size of the meltpool is one of the most important parameters in laser metal deposi- tion.

The melt pool size is highly dependant on the distance between the focussing lens and workpiece which inevitably changes during the laser metal deposition process.Hence, monitor- ing the size of the melt pool is an important part of the quality control process.Conventional image vision processing techniques use infrared ﬁlters along with a high speed shutter along with synchronising the laser with the shutter[4]. These propietary vision systems are relatively expensive to buy and maintain. Previous approaches to image the meltpool include [3]1. Filtering out IR radiation along with pinhole concept2.

These propietary vision systems are relatively expensive to buy and maintain. Previous approaches to image the meltpool include [3]1. Filtering out IR radiation along with pinhole concept2. Dichroic ﬁlters with polarization lensBoth the previous approaches were unsuccessful due to the the inability of the lens to ﬁlter out the large amount of IR radiation emitted by the melt pool.We try to implement a lowcost alternative solution to the conventional image processing process which enables us to determine the size of the melt pool and study the eﬀect of various system parameters on the melt pool size.3 ConceptDuring the laser deposition process, the laser heats up a ﬁnite area of the substrate and forms a meltpool.This meltpool radiates out energy of diﬀerent wavelengths depending on the type of laser used,intensity,substrate material etc. According the black body radiation theory ,the wavelength of light emitted by the melt pool cannot be in the ultraviolet range.

According the black body radiation theory ,the wavelength of light emitted by the melt pool cannot be in the ultraviolet range. According to Aalderink et al,[5] the external illumination concept can be utilized if the signal to noise ration is of suﬃcient qualityS/N = IS/INwhere IS is the irradiance of reﬂected illumination light that reaches the image sensor and IN is the irradiance of the welding process that reaches the image sensor.

We seek to eliminate the infrared radiation emanating from the melt pool entirely while providing our external illumination of diﬀerent wavelength onto the melt pool.Figure 1: The CIE 1931 XYZ Color Space relates black body radiation color and tempera- ture.Image courtesy of Wikimedia CommonsExternal illumination in the form of highly intense ultraviolet light ,when incident uponthe meltpool should provide a clear image using special ﬁlters.4 Experimental setupThe MicroLAMP (µ LAMP) consists of a 50 Watt(W) ﬁber coupled diode laser whose fo- cussing lens is mounted on the vertical Z axis of a CNC system while the substrate is placed on the X-Y motion table. The focussing lens of laser is maintained at a constant standoﬀ dis- tance of 54 mm while the substrate is moved beneath the laser. The substrates were cleaned prior to deposition using isopropyl alcohol to remove any dirt or grease on the surface.

The substrates were cleaned prior to deposition using isopropyl alcohol to remove any dirt or grease on the surface. The CNC table is driven by an Sanguino microcontroller connected to a computer via USB. TheFigure 2: Experimental setup of Laser and vision systemsubstrate is secured in a vise and an inert atmosphere of argon gas is provided to prevent oxidation of the deposit on the substrate.4.1 Image Processing setup4.2 External illuminationAn array of Ultraviolet(UV) light emitting diodes were focussed onto the melt pool area. The wavelength of ultraviolet light emitted is in a very speciﬁc range of 390-395 nm. The LED’s are run at 90 % of their maximum capacity to avoid burning out and ensure longer operating life. Since the peak spectral response of the LED’s are at 18 degrees emission angle , the UV array is tilted appropriately towards the substrate to ensure that the melt pool receives the highest spectral intensity of UV light.

A ﬁxture was designed and prototyped by a Fused deposition Modeller to hold the LED’s along with the USB microscope.The total optical power output of the array is approximately 52.92 mW5Image acquisition setupA Dinolite USB microscope with a CMOS sensor was selected with a maximum magniﬁcation of 200X . A CMOS sensor was preferred since it has high sensitivity in the ultraviolet range compared to CCD sensors. Hence the reﬂected UV radiation coming oﬀ the meltpool will be absorbed by the image sensor present in the microscope.The microscope was ﬁtted with a narrow band pass ﬁlter which only allows UV light of wavelength 389-399 nm to pass through.A traditional Infra Red (IR) ﬁlter is used along with the narrow band pass ﬁlter to remove any unwanted IR from the image.

The white LED illumination lights inbuilt within the microscope were removed due to interference with the image.The microscope and UV lights are mounted in a rapid prototyped ﬁxture by a Dimension Fused Deposition Modeller. The ﬁxture was designed as to allow various angular adjustements to the ultraviolet illumination and the microscope.Figure 3: CMOS sensitivity to diﬀerent wavelengths6 PostprocessingThe raw image from the USB microscope is post processed using OpenCV image pro- cessing libraries in Python. The various post processing steps carried out include ﬁlter- ing,thresholding,smoothing. OpenCv software was used because of its ease of use in image processing and also because it was open source .Since the microscope is mounted at an angle to the meltpool perspective correction is applied to remove perspective error. The various post processes carried out include1.

The various post processes carried out include1. Obtaining the image-The output from the usb microscope is video.A frame is grabbed from the video to do image processing2. Converting the image to grayscale-The intensity values of light are observed in a grayscale image3. Smoothing the image-To remove noise in the image4. Thresholding the image-This is done to eﬀectively diﬀerentiate the melt pool from the background.5. Contour ﬁtting and ellipse ﬁtting-Since the melt pool obtained is not always circular,an ellipse is ﬁtted to the melt pool contour to obtain an average measurement of the melt pool size6.

Contour ﬁtting and ellipse ﬁtting-Since the melt pool obtained is not always circular,an ellipse is ﬁtted to the melt pool contour to obtain an average measurement of the melt pool size6. Measuring melt pool sizeThe necessary camera intrinsic matrix and distortion coeﬃcients are determined using inbuilt calibration functions in OpenCv which enables us to obtain the best melt pool size ap- proximation and minimize radial,tangential distortions [6].Since the microscope is not mounted coaxially with the laser,it is necessary for us to obtain a bird’s eye view of the melt pool to avoid error.Figure 4: UV array and microscope setup7 ResultsThe following image processing algorithm was developed using OpenCv image processing library.The raw image obtained from the vision system is subjected to the above image pro- cessing to obtain the melt pool size. Other parameters such as circularity can be easily ascertained from the data.

Other parameters such as circularity can be easily ascertained from the data. The overall cost of the system , which includes cost of the usb microscope, ﬁlters and LED’s is approximately 400 dollars excluding the cost of ﬁxturing which is relatively inexpensive compared to commercial systems available in the market today.Figure 5: Raw image obtained from camera8 Conclusion and Future WorkThe vision system is currently capable of acquiring images of the melt pool and processing it to acquire useful data,speciﬁcally melt pool size. Due to equipment problems we were unable to obtain suﬃcient real world data from the vision system.Future work involves integrating the imaging system with the mini laser deposition system so that in process changes are detected and corrected to obtain repeatable and good quality deposits.Figure 6: ellipse ﬁtted onto melt pool9 AcknowledgementThis research was supported by the National Science Foundation grants IIP-0637796, and IIP- 0822739.

Support from Product Innovation and Engineering, LLC, Missouri ST Intelligent Systems Center, and the Missouri ST Manufacturing Engineering Program, are also greatly appreciated.ReferencesEhsan Toyserkani and Amir Khajepour and Stephen Corbin , Laser Cladding, CRC press, 2nd Edition, 2005W. Liou, J. Choi, R. G. Landers, V. Janardhan*, S. N. Balakrishnan*, and S. Agarwal , Research and Development of A Hybrid Rapid Manufacturing Process, University of Missouri RollaTodd Sparks and Lie Tang and Frank Liou, Development of a Melt Pool Tracking Vision System for Laser Deposition, Department of Mechanical and Aerospace Engineering, Missouri University of Science TechnologyYoshimi Kizaki and Hirozumi Azumi and Satoshi Yamazaki and Hisashi Sugimoto and Soya Takagi, Phenomenological studies in Laser Cladding.Part2.Thermometrical Experiments on the meltpool , Journal of Applied Physics Vol(32) 1993B.J. Aalderink , R.G.K.M. Aarts , J.B. Jonker and J.

viewed PaperDEVELOPMENT OF A LOW THERMAL EXPANSION SLA RESIN FOR NICKEL PLATING APPLICATIONSC. Hammond*, J. Greenstreet*, W. Gomez*, R. Dang*, and J. Tate*†Ingram School of Engineering † Materials Science Engineering and Commercialization Program †Email: JT31@txstate.edu Phone: (512) 245-1826AbstractStereolithography additive manufacturing is a method of producing parts by stacking layers of a photopolymer resin cured by exposure to UV light . This method of additive manufacturing gives great resolution, but often lacks the material properties of other techniques. One method to increase part performance is the addition of a thin nickel plating to increase strength, heat deflection, and chemical resistance. No solution has been proposed for using nickel-coated parts in harsh environments where the large difference in thermal expansion rates between the nickel plating and base resin cause internal stresses to form.

The excellent chemical resistance of the nickel plating would also allow these parts to be used in high-temperature, oxygen-rich environments such as those presented in life support systems. Through the development of a high- performance nanocomposite SLA resin, we hope to achieve parts with good mechanical properties and a CTE similar to nickel.IntroductionStereolithography (SLA) 3D printing is a method of additive manufacturing that involves selectively curing thin layers of a photosensitive resin, usually by ultraviolet radiation. By curing one thin cross-section (usually around 50 microns) at a time, a complex part can be fabricated with features that could not be produced using traditional subtractive manufacturing. This technique results in accurate parts with good surface finish and can be achieved through relatively low-cost commercial off-the-shelf (COTS) machines.

This technique results in accurate parts with good surface finish and can be achieved through relatively low-cost commercial off-the-shelf (COTS) machines. This method also allows for rapid iteration with the ability to quickly export and print a 3D model with minimal setup time. The drawback of SLA additive manufacturing is still widely constrained by material properties of the photopolymers. Research has been performed on improving the properties of these photopolymers through the addition of different plasticizers and chemistries. Even with these innovations, SLA parts cannot match the strength, chemical resistance, or heat deflection of metal or high-performance fused deposition modeling (FDM) parts.

Even with these innovations, SLA parts cannot match the strength, chemical resistance, or heat deflection of metal or high-performance fused deposition modeling (FDM) parts. In this paper, we will be attempting to develop a low thermal expansion, high strength resin that can be nickel-plated for increased heat and chemical resistance.Companies such as Formlabs have experimented with nickel-coated SLA parts, withsuccess for low-temperature applications [1].162Figure 1: Nickel-plated SLA brackets by Formlabs [1]Although large increases in strength and stiffness at room temperature were recorded, Formlabs recommends that the parts not exceed 150ºC, or be subjected to thermal cycling. If a resin could be formulated with a CTE matching the metal coating, nickel-coated SLA parts could find new use cases in high-temperature applications.Literature reviewVarious research has been performed on improving the mechanical properties of SLA resins.

Sano and Matsuzaki in 2018 [2] experimented with printing composite materials by adding in ground glass powder, short length fibers, and continuous fiber. The results showed large improvements in mechanical properties for every fiber type.Figure 2: Mechanical properties of glass powder and fiber filled SLA prints [2]Plasticizers or fillers can also be added to the resin to improve performance. Nanoparticles of silver, silica, copper, and other materials have been used in research to improve part strength, electrical, and thermal conductivity [3]. Very low and even negative coefficient of thermal163expansions (CTE’s) have been achieved in aligned multiwalled carbon nanotube thin-film composites [4]. However, there is a practical limit to the amount of filler you can add to a resin. In most cases, adding more than 10% by weight of filler can block ultraviolet light from reacting with the resin, causing print failures.

In most cases, adding more than 10% by weight of filler can block ultraviolet light from reacting with the resin, causing print failures. Heavy particles such as copper and iron can also begin settling out of the resin and will result in uneven dispersion which can cause stress concentrations within the part and non-isotropic properties which are undesirable. Using nanoparticles of materials with low thermal expansion rates, the coefficient of thermal expansion (CTE) of the resin can be matched as that of the nickel. These nanoparticles will also have to be around the same density as the resin to facilitate good dispersion and reduce settling during printing. The main materials of interest are multiwalled carbon nanotubes (MWCNT’s), silicone dioxide nanoparticles, and boron nitride nanoparticles. These additives were chosen for their low coefficient of thermal expansion (CTE) and high young’s modulus which are important in reducing the composites overall thermal expansion rate.

These additives were chosen for their low coefficient of thermal expansion (CTE) and high young’s modulus which are important in reducing the composites overall thermal expansion rate. Carbon nanotubes with a mean diameter of 41 nanometers have been shown to have a negative CTE value between 0º and 100˚ C , roughly -1 µm/m°C [5]. The goal is to reduce the CTE of the resin down to 13.3 µm/m°C as that of nickel’s while using the least amount of filler possible.Figure 3: CTE (K-1) comparison of carbon nanotubes [5]EquipmentThe printer chosen for experimentation is a mUVe3d LCD 2k pro for its open-source design and high output UV lamp. The UV LED array can output up to 82 watts of 405 nm UV light across its 5.5 inch 2k resolution LCD display.164Figure 4: mUVe LCD 2k Pro printerSLA resins for 3D printing are composed of three main constituents, an oligomer, monomer, and photoinitiator.

Oligomers chosen for our formulation include CN154 epoxy methacrylate and Ebecryl 3700 BPA Epoxy resin. These epoxies were chosen for their high hardness and temperature resistance. Monomers, sometimes referred to as reactive diluents, are added to the mixture in order to reduce the viscosity down to a level where it can be easily poured into the resin vat. A low viscosity is also essential as the mixture must reflow under the print bed after every layer, poor flow in a resin can result in voids in the final print. Our reactive diluent is Trimethylolpropane triacrylate, a trifunctional acrylate ester monomer used for its fast cure response and high crosslinking capability.

Finally, photoinitiator are used to sensitize the resin to UV light and start the polymerization process.Figure 5: UV resin polymerization [6]165Table 1: Chemicals used in resin formulationsCHEMICAL NAME CN154 BPA EPOXY METHACRYLATE EBECRYL 3700 BPA EPOXY RESIN TRIMETHYLOLPROPANE TRIACRYLATE (TMPTA) ACRYLATE ESTER DIPHENYL(2,4,6- TRIMETHYLBENZOYL)PHOSPHINE OXIDE (TPO-L) PHENYLBIS(2,4,6- TRIMETHYLBENZOYL)PHOSPHINE OXIDE (BAPO) BYK-W 9010USE OligomerOligomer Monomer / reactive diluentphotoinitiatorphotoinitiatorSUPPLIER Sartomer AmericasAllnex AllnexSigma-AldrichSigma-AldrichDispersing AgentBYKNanoparticle additives include boron nitride, silicon dioxide, and multi-walled carbon nanotubes.Table 2: Nanoparticle and average particle sizeNanoparticleAverage Particle sizeMulti-Walled carbon nanotubes30-50nm outer diameterBoron Nitride60-70nmSilicon Dioxide80nmIncorporation of nanoparticles into the resin was achieved through sonication, a well understood technique for nanoparticle dispersion [7].

This is a method where a high frequency transducer is used to create pressure waves in the resin causing cavitation to occur, releasing large amounts of energy into the mixture and dispersing the clustered nanoparticles.ExperimentationIn order to get a baseline resin standard that we can compare to the nanocompositeformulations, a neat resin was created with both of the epoxies.166Table 3: Ebecryl based Neat resin formulationChemicalAmountEBECRYL 3700 BPA epoxy resin40% by weight (20g)TMPTA58% by weight (29g)BAPO2% by weight (1g)To reduce the high viscosity of the epoxy resin and make incorporation easier, the Ebecryl resin was heated in a 70ºc water bath. It was then weighed and incorporated into the TMPTA and photoinitiator solution and mixed by hand for approximately 5 minutes. The mixture was then placed under a partial vacuum (-27 inHg) for degassing. Initial test prints yielded good results with great accuracy at a layer exposure time of 8 seconds.

Initial test prints yielded good results with great accuracy at a layer exposure time of 8 seconds. We did see some small defects from air entrapment in the mixture, possible due to either inadequate degassing, or air incorporation during the printing process.Figure 6: Ebecryl Neat resin test printA baseline resin using CN154 epoxy was also tested for comparison, this formulation wasprocessed identically to the Ebecryl mixture.167Table 4: CN154 based Neat resin formulationChemicalAmountCN154 epoxy resin40% by weight (20g)TMPTA58% by weight (29g)BAPO2% by weight (1g)Table 5: Neat resin print parametersPrint settings Layer height(μm)Cure time (Seconds)Pause time before print (Seconds)Bed lift speed (μm / Second)Burn-in layers (5)50101200Standard layers5080.15200Nanocomposite mixtures were then created by adding 1 gram of BYK-W 9010 wetting and dispersing agent to the neat resin mixtures along with the nanoparticles.

Additionally, nanocomposite mixtures received 1.5 grams of photoinitiator compared to the 1 gram in the neat resin mixtures to ensure sufficient polymerization. For composite mixtures containing multi- walled carbon nanotubes (MWCNT’s), an additional 0.5g of TPO-L initiator was incorporated into the resin. The mixtures were then stirred by hand for initial incorporation before being sonicated in in a cold water bath. After every 6 sonication cycles, the process was paused and the resin was hand mixed for approximately 30 seconds. This was done to ensure even heat dissipation and to prevent pockets of resin prematurely polymerizing from excess localized heat.

This was done to ensure even heat dissipation and to prevent pockets of resin prematurely polymerizing from excess localized heat. After sonication, the composite resin mixture was degassed under vacuum.Table 6: Sonication parametersNumber of sonication cycles Cycle on time Cycle off time Power24 10 seconds 20 seconds 40% (30 Watts)168Figure 7: Incorporation, sonication, and printing of nanocompositesPrint parameters for the nanocomposites can be found below in tables 7, 8, and 9. Nanoparticle fillers block ultraviolet light from reacting with the resin matrix and thus require extended cure times versus the neat resin samples.

CN154 samples were noticeably more optically transparent coming off the printer than the Ebecryl samples. After 5 minutes of UV post cure, both resins became tinted yellow.Figure 8: Post print cure, post cure yellowingPrints containing boron nitride and carbon nanotubes showed large changes in the appearance of the sample. We also observed a small amount of flashing on the bottom of the carbon samples, this is likely from overexposing the initial burn-in layers, causing excess curing outside of the masked area. The boron test samples also grew slightly in all dimensions compared to the other fillers, we suspect this is from UV light reflecting off the white boron particles and causing sample expansion.170Figure 9: Boron nitride, carbon nanotube samplesPrints containing silicone dioxide looked very similar to the neat resin samples despite the nanoparticles having a bright white color, similar to the boron nitride before incorporation.

This is possibly from the resin and sio2 having very similar refractive index’s [8].Figure 10: Sio2 sample on print bedOverall, we demonstrated the ability to successfully print nanocomposite resin systems on a COTS SLA printer. Neat resin samples demonstrated very similar exposure times and processing methods to commercially available resins. Nanoparticle composites do require higher exposure171times and extra processing steps such as sonication and vacuum degassing, but still demonstrated reliable printing.Results and DiscussionTensile testing was performed using ASTM D638 type IV dog bone samples at 1mm/minute on an MTS Exceed universal test system. The slower testing speed was chosen due to the brittle nature of the epoxy resin samples. With the samples being too small for use with an extensometer, sample strain was estimated as a function of crosshead movement.

With the samples being too small for use with an extensometer, sample strain was estimated as a function of crosshead movement. All samples were tested with the layer lines running parallel to the axis of measurement..Figure 11: Tensile testing setupFor a baseline standard, the neat resins created with CN154 and Ebecryl were compared tothe commercial Stratasys “Digital ABS plus”.172Figure 12: Tensile results of neat resinsTable 10: Neat resin tensile testingResinUltimate tensile strengthPercent elongationYoungs ModulusStratasys Digital ABS Plus CN154 Ebecryl5.68 Mpa17.56 Mpa 8.55 Mpa5.3%0.74% 0.73%93.3 ± 1.53 MPa2557.7 ± 1.4 MPa 1241.1± 0.4 MPaAs shown in figure 12, both of the epoxy resins have greater strength and stiffness compared to the commercial Stratasys “Digital ABS Plus” UV resin. CN154 also exhibited a much higher ultimate strength, breaking at 346 newtons compared to the 112 newtons for the commercial system.

CN154 also exhibited a much higher ultimate strength, breaking at 346 newtons compared to the 112 newtons for the commercial system. The combination of a high performance epoxy with the tri-functional monomer results in a stiff, but brittle part.With the addition of nanoparticles to the CN154 based resin, silica and carbon nanotubes showed increased elongation, but reduced the ultimate strength. Samples containing 1gram of boron nitride showed both decreased strength and elongation.173Figure 13: CN154 nanocomposite tensile testingTable 11: CN154 nanocomposite tensile resultsResinCN154 Neat CN154 2%wt BN CN154 2%wt Sio2 CN154 0.2%wt MWCNTUltimate tensile strength 17.56 MPa 6.65 MPa 8.45 MPa12.81 MPaPercent elongation 0.74% 0.54% 0.93%0.77%Youngs Modulus2557.7 ± 1.4 MPa 2432.9±1 MPa 1029.2±0.2 MPa1763.1±0.3 MPaWe believe the slight decrease in performance is due to poor distribution of the boron and carbon in the resin.

Over the long print times required, it was observed that clumps of carbon nanoparticles began to conglomerate on the print bed. Poor distribution could cause stress concentrations within the test samples, thus resulting in a decrease in mechanical properties. We also believe that we had poor adhesion between the nanoparticles and resin because samples identically processed using the Ebecryl epoxy resin showed large gains in mechanical properties with the addition of carbon nanotubes.174Figure 14: Carbon nanotube conglomeration on print bedEbecryl based silica and MWCNT nanocomposites showed large strength and elongation gains over the neat resin.

Samples containing 1gram of boron nitride exhibited almost identical results to the neat resin samples.Figure 15: Ebecryl Nanocomposites tensile results175Table 12: Ebecryl nanocomposite tensile resultsResinEbecryl Neat Ebecryl 2%wt BN Ebecryl 2%wt Si Ebecryl 0.2%wt MWCNTUltimate tensile strength 8.55 MPa 8.55 Mpa 16.81 Mpa17.47 MpaPercent elongation 0.73% 0.72% 1.0%0.81%Youngs Modulus1241.1± 0.4 MPa 1231.8±0.9 MPa 1895.9±0.9 MPa2319±1.1 MPaThermal characterization was performed on a Waters Corporation TA Instruments TMA Q400 using a standard atmosphere from ambient temperature to 150˚ Celsius. CTE was measured by performing an alpha fit operation to find the slope of sample expansion vs temperature in TA Universal Analysis from 60˚ to 140˚C.Figure 16: CTE sample puck loaded into TMA Q400Baseline resin samples using Stratasys “DigitalABSPlus” UV resin were printed at a 50- micron layer height and 8 second cure time per layer.

TMA measurements showed an average CTE of 203.2 µm/m(m.ºC) from 60-140 ºC in a standard atmosphere for samples with the layer lines running perpendicular to the axis of measurement. Samples with the layer lines parallel to the axis of measurement recorded a 1.5% difference in thermal expansion at 206 µm/m(m.ºC). From this small difference, we can conclude that the unmodified resin samples are thermally176isotropic. All CTE samples tested were printed as 5x6mm cylindrical pucks with the layers running perpendicular to the axis of measurement. The epoxy-based resins also showed increased performance in CTE analysis.

The epoxy-based resins also showed increased performance in CTE analysis. At 140˙C we did observe sample cracking with the CN154 resin from thermal stress.Figure 17: Neat resin CTE resultsTable 13: Neat resin CTE resultsResin Stratasys Digital ABS Plus Ebecryl CN154CTE 203.2 µm/m(m.ºC) 164.9 µm/m(m.ºC) 126.4 µm/m(m.ºC)Percent change Baseline measurement 18.84% reduction 37.79% reductionCN154 nanocomposites eliminated the stress fracture seen in the neat resin sample. Both the carbon nanotubes and boron nitride samples had an increased CTE compared to the neat resin.

Both the carbon nanotubes and boron nitride samples had an increased CTE compared to the neat resin. With 1 gram of silica, a 5.69% reduction in CTE was measured.177Figure 18: CN154 composite CTE resultsTable 14: CN154 composite CTE resultsResin CN154 Neat resin 0.2% wt MWCNT 2% wt BN 2% wt SiO2CTE 126.4 µm/m(m.ºC) 148.2 µm/m(m.ºC) 158.6 µm/m(m.ºC) 119.2 µm/m(m.ºC)Percent change Baseline measurement 17.24% increase 25.47% increase 5.69% reductionCTE results for the Ebecryl based samples showed similar trends to the CN154 samples. Both boron nitride and MWCNT resulted in increased thermal expansion while silica reduced the composite CTE by 7.82%. We observed sample failure at 110ºC in the MWCNT composite with large cracks forming on the surface.

We observed sample failure at 110ºC in the MWCNT composite with large cracks forming on the surface. To account for the large failure, CTE was measured from 40- 100ºC on the carbon sample only.178Figure 19: Ebecryl composite CTE resultsTable 15: Ebecryl CTE results (*Indicates changed measurement parameters)Resin Ebecryl Neat resin 0.2 % wt MWCNT 2% wt BN 2% wt SiO2CTE 164.9 µm/m(m.ºC) 174.5 µm/m(m.ºC)* 182.5 µm/m(m.ºC) 152.0 µm/m(m.ºC)Percent change Baseline measurement 5.82% increase* 10.67% increase 7.82% reductionComparison of all CTE measurements can be found below in table 16.

This is due to the air bubbles that have been occurring during the mixing process of both mixtures of resins. After a direct comparison of both surfaces of each print we notice that there are significantly more defects in the MWCNT prints compared to the Boron Nitride prints. As carbon is an effective absorber of UV light [9], we concluded that the additional defects were caused from incomplete polymerization during the printing process.Figure 20: Optical microscope images of MWCNT prints and its surface defectsFigure 21: Optical microscope images of Boron Nitride prints and its surface defects180ConclusionThis research demonstrated the ability to modify the mechanical and thermal properties of SLA resin through the addition of nanoparticles. For CTE reduction, silica was identified as the best performing filler, reducing thermal expansion in both epoxy resin formulations.

For CTE reduction, silica was identified as the best performing filler, reducing thermal expansion in both epoxy resin formulations. Tensile performance was also improved in the Ebecryl formulation and increased elongation was observed in the CN154 resin with the addition of silica. Sio2 is a very popular additive in epoxy as it demonstrates strong adhesion to the resin matrix as long as good distribution is achieved. This great matrix adhesion explains the increased thermal characteristics in both the Ebecryl and CN154 samples. Multi-walled carbon nanotubes showed great improvements in mechanical testing for the Ebecryl based resin, but did not deliver the same benefits in the CN154 composite. As both mixtures were processed identically, we believe this difference is due to the increased wetting and bonding between the nanoparticles and Ebecryl resin matrix.

As both mixtures were processed identically, we believe this difference is due to the increased wetting and bonding between the nanoparticles and Ebecryl resin matrix. Even with the addition of BYK-W 9010 wetting and dispersing agent, the CN154 resin failed to properly bond with the nanotubes, resulting in decreased performance. Carbon nanotubes have also been shown to increase the thermal conductance of composite systems [10], a possible explanation for the increase in CTE for both resin systems. This would also explain why we saw sample failures and cracking occur during the thermal testing of carbon samples. Boron nitride nanoparticles failed to increase mechanical or thermal performance in either resin matrix. This probably due to a couple of factors, ceramics such as boron nitride have been shown to decrease cure depth substantially in SLA resins [11]. Decreased cure depth could cause insufficient polymerization during the print, resulting in weaker parts.

Decreased cure depth could cause insufficient polymerization during the print, resulting in weaker parts. We could also be experiencing the same wetting problem seen with the carbon nanotubes, as our wetting agent is optimized for carbon fillers and might not be effective in wetting and dispersing ceramics. Boron nitride is also known for being difficult to properly disperse [12]. Poor dispersion of the boron could be from the limited power output of our sonication machine, or could be occurring during the printing process as the low viscosity resin allows the nanoparticles to conglomerate and settle out during the long print times. In total, a combination of these factors could be the cause of the boron nitride being an ineffective filler for SLA additive manufacturing.1811.2.3.4.5.6.7.8.9.10.11.12.ReferencesFormlabs, Strength from Metal: Stratagies and use cases for Electroplating SLA Parts. 2021.

velopment of a Low-cost Parallel Kinematic Machine for Multi-directional Additive ManufacturingXuan SongYayue PanYong Chen*Daniel J. Epstein Department of Industrial and Systems Engineering University of Southern California, Los Angeles, CA 90089, U.S.A.Corresponding author, Phone: (213) 740-7829, Fax: (213) 740-1120, Email: yongchen@usc.eduABSTRACTMost additive manufacturing (AM) processes are layer-based with three linear motions in the X, Y and Z axes. However, there are drawbacks associated with such limited motions, e.g. non-conformal material properties, stair-stepping effect, and limitations on building-around-inserts. Such drawbacks will limit additive manufacturing to be used in more general applications. To enable 6-axis motions between a tool and a work piece, we investigate a Stewart mechanism and developed a low-cost prototype system for multi-directional additive manufacturing processes such as the Fused Deposition Modeling (FDM) and CNC Accumulation.

The technical challenges in developing such an AM system are discussed including the hardware design, motion planning and modeling, platform constraint checking, tool motion simulation, and platform calibration. Several test cases are performed to illustrate the capability of the developed multi-directional additive manufacturing system.KEYWORDS: Additive manufacturing, multi-direction, parallel kinematic machine, fused deposition modeling, building-around-inserts.1 IntroductionAdditive manufacturing (AM) processes can directly fabricate three-dimensional (3D) computer- aided design (CAD) models by controlling the selective accumulation of materials. Most AM processes are layer-based, that is, a given 3D model is first sliced into a set of two-dimensional (2D) layers; accordingly, a physical part is fabricated by stacking the sliced 2D layers together to approximate the given CAD model. An example of a tilted rod (AB) is shown in Figure 1.

An example of a tilted rod (AB) is shown in Figure 1. For such layer-based AM processes, only the linear motions in the X, Y and Z axes are required.Figure 1: A schematic illustration of additive manufacturing process.The layer-based fabrication approach has many benefits. For example, (1) tool path planning and hardware design are simplified; and (2) complex shapes that are impossible to be made before can be fabricated. However, there are also drawbacks associated with the layer-based fabrication approach. For example, (1) the surface finish is poor due to the stair stepping effect (refer to rod AB in Figure 1). (2) Also the material property of a geometric feature will depend on the building direction that is used in the fabrication process. Consequently, the material property of a tilted rod in different tilting angles will be different. In addition, (3) it would be difficult to build parts around inserts (e.g.

Consequently, the material property of a tilted rod in different tilting angles will be different. In addition, (3) it would be difficult to build parts around inserts (e.g. an electric or optical component) due to the limited tool motions that are allowed in the system.To address the problems of the layer-based AM processes, various methods have been proposed. For example, controlled cure depth [2, 3], post-processing [4, 5] and meniscus methods [6] have been developed for improving surface finish; and techniques such as model shape modification [7] and hybrid process development [8] have been employed to enhance the fabrication capability of building-around- inserts. However, most approaches can only improve one or a few drawbacks in a limited fashion mainly due to the use of a single build direction (Z axis) and a uniform layer thickness in the building process.

However, most approaches can only improve one or a few drawbacks in a limited fashion mainly due to the use of a single build direction (Z axis) and a uniform layer thickness in the building process. In comparison, multi-directional AM processes, in which materials are added along multiple directions using non-uniform layer thickness (refer to Figure 2), can address the limitations of the layer-based AM processes. In the future AM process development, both layer-based and non-layer-based fabrication approaches may be required in the material deposition process.Figure 2: Comparison between the layer-based and multi-directional AM processes.To achieve multi-axis motion between the accumulative tool and the workpiece, multi-directionalAM processes can be classified into (1) platform based and (2) accumulative tool based approaches.(1) Platform based approaches.

Multi-directional fabrication can be achieved by rotating the platform in order to orientate the workpiece related to the accumulative tool. For example, a customized compliant parallel kinematic machine is presented in [9] for the Multi-Direction Layered Deposition (MDLD) process. The machine is comprised of two parts: a X-Y overhang head unit, and a workpiece orientation unit to rotate the built part to have a building direction that is aligned to the head unit. Consequently the deposition head can add materials from different orientations. The Laser Direct Casting (LDC) process [10] uses a similar approach by rotating built parts to achieve multi-directional fabrication. A multi-directional UV lithography process at micro- and nano- scales is also discussed in [11], in which two step motors are used to control the tilting and rotational angles of the substrate.

A multi-directional UV lithography process at micro- and nano- scales is also discussed in [11], in which two step motors are used to control the tilting and rotational angles of the substrate. In the design of a laser direct metal deposition system [12] for repairing deep and internal cracks in metallic components, the laser beam is kept stationary while the workpiece is moved and rotated.(2) Accumulative tool based approaches. In multi-directional AM processes, the accumulative tools, instead of the platform, can be oriented with multiple degrees of freedom. A multi-directional Direct Metal Deposition (DMD) system using a high-power laser is presented in [13-16]. The laser beam is focused onto a workpiece and produces a melt pool. Medal powders are injected into the melt pool by feeding with inert gas stream. In the system, the laser head is controlled by a 5-axis motion mechanism that allows deposition of given shapes.

Medal powders are injected into the melt pool by feeding with inert gas stream. In the system, the laser head is controlled by a 5-axis motion mechanism that allows deposition of given shapes. As shown in [16], the slicing direction can be arbitrary, and a number of layers with non-uniform thicknesses can be generated. Accordingly, the tool path planningsoftware can convert CAD models into nozzle motions for multi-axis deposition. The Directed Light Fabrication (DLF) process [17] is another kind of direct metal deposition process that fuses inert-gas- delivered metal powders into the focal zone of a high powered laser beam. By using the multi-axis numerical control sequence tool paths, materials can also be deposited in varying orientations.

By using the multi-axis numerical control sequence tool paths, materials can also be deposited in varying orientations. For the fabrication of polymer-based objects, a Computer Numerically Controlled Accumulation (CNCA) process was developed [18], in which a fiber optical cable connected with a high power UV-LED is controlled by a 5-axis motion system to enable the X, Y, and Z axes translations and the A and B axes rotations. Therefore, the accumulative tool can cure liquid resin in various directions.Compared with the tool based approaches, the platform based processes have simpler structures, especially when the system only needs a few degrees of freedom (DOFs). However, rotating the platform is less flexible with limited degrees of freedom. To enable 6-axis motions between a tool and a work piece, a Stewart mechanism is investigated in this paper for developing a low-cost multi-directional additive manufacturing system.

To enable 6-axis motions between a tool and a work piece, a Stewart mechanism is investigated in this paper for developing a low-cost multi-directional additive manufacturing system. Compared with the traditional translation and rotation based approach, the Stewart mechanism enables the system to be less bulky. Two kinds of accumulative tools including a FDM heating extruder and a fiber-optics-based CNC accumulation tool are considered. The rest of this paper is organized as follows. Section 2 introduces the hardware design based on the Stewart mechanism. The cost of the prototype system is also discussed. In section 3, a kinematic modeling and simulation software system is presented. Section 4 describes a calibration method to achieve improved accuracy of the parallel kinematic machine. Section 5 presents the data processing pipeline of the multi-direction AM system based on the machine.

Section 5 presents the data processing pipeline of the multi-direction AM system based on the machine. Section 6 demonstrates the part fabrication of the prototype AM system using a FDM heating extruder. Finally conclusions are drawn in section 7.2 Hardware DesignThe hardware components of our parallel kinematic machine are introduced in this section. To enable the 6-DoF motion of a tool with respect to a fixed frame, six length-changeable struts are used to connect a moving platform on which an accumulative tool is mounted. For each strut, one of its ends is connected to the moving platform by a 3-DoF joint, and another end is connected to a fixed base frame by a 2-DoF joint (refer to Figure 3). This 6-axis parallel kinematic machine is also called Gough-Stewart mechanism. In industry, this mechanism has been used in precision positioning system (e.g. Hexapod 6- axis parallel positioning systems from Physik Instrumente).

In industry, this mechanism has been used in precision positioning system (e.g. Hexapod 6- axis parallel positioning systems from Physik Instrumente). However, the commercial systems are expensive with relatively small travel ranges (<50mm).Figure 3: Schematic illustration of a Stewart mechanism.Motivated by the recent progresses on developing low-cost 3D printers (e.g. Makerbot, Solidoodle, etc.), we investigated the development of a low-cost 6-axis parallel positioning system for additive manufacturing processes. As shown in Figure 3, the parallel kinematic machine is made up of six length- changeable struts. In our testbed, six ball screw linear actuators are used as the length-changeable struts. The linear actuators are connected in pairs to a hexagon moving platform by ball joints. Each motor body is mounted on a customized universal joint, which has 2-DoF motion relative to the fixed base frame.The CAD model of the designed system is shown in Figure 4.

Each motor body is mounted on a customized universal joint, which has 2-DoF motion relative to the fixed base frame.The CAD model of the designed system is shown in Figure 4. In our testbed, we used linear actuators from Eastern Air Devices Inc. (Dover, NH). Its original lead-screw was replaced by a 1/4-16 ACME one with a length of 12”. The system is controlled by a high performance 8-axis motion control board KFLOP+2KSTEP (Dynomotion Inc., Calabasas, CA). A motion parameter generation and control software system has been developed. The system can load in the G-code of tool paths, transform them into motion command parameters, and send them to the motion controller though 6 output pins in the KSTEP board. A photo of the built prototype system is shown in Figure 11. Based on the Stewart mechanism, relatively small motions of the linear actuators can lead to large motions of the tool on the moving platform.

Based on the Stewart mechanism, relatively small motions of the linear actuators can lead to large motions of the tool on the moving platform. Hence the multi-directional AM system can have high fabrication speed.Figure 4: CAD model design of our parallel kinematic machine.Figure 5: The design of a length-changeable strut.The detail design of a length-changeable strut based on a linear motor is further shown in Figure 5. For the linear actuator, it creates linear motions only when the end of the lead screw has a holding torque. To impose the required torque, we use a pair of slide bushings with two guide rails on each motor. Our design has good modularity since all the struts in the designed system are exactly the same.Table 1 is the expense of the components that are used in the prototype system. The total cost is less than $1,500 excluding the base frame.

The total cost is less than $1,500 excluding the base frame. Compared to the commercial hexapod machines, it has a rather low cost and has the potential to be used in developing low-cost 3D printers with multi-directional motions. The 6-axis motion system is general. Various accumulative tools can easily be integrated in the system. Both the FDM and CNCA processes are demonstrated using the developed prototype system.

The 6-axis motion system is general. Various accumulative tools can easily be integrated in the system. Both the FDM and CNCA processes are demonstrated using the developed prototype system. In addition, since all the motions are conducted by moving the accumulative tools, the system can have low inertia with small energy consumption.Table 1: Cost of the prototype systemItemQuantityPrice eachLinear actuator6$75Aluminum plate4"×4", 12"×24"$78Aluminum angle4 '$32Bearing24$7.0Lead screw6$103"alloy steel thread stud12$1.5Ball joint6$7.0Aluminum rod Others (screw, nuts, slide, bushing, etc.) Motor controller Total6'NA1 NA$10$80$500 $1,4383 Motion Planning and Software System DesignThe desired tool motion in the 6-axis parallel kinematic machine needs to be converted into the linear motions of the six length-changeable struts.

One coordinate system is attached to the fixed base frame (B) and another one is fixed on the moving frame (P). Initially, the moving frame has the same coordinate axes as those of the base frame (refer to OP and OB in the figure). During the fabricationprocess, the coordinate system of the moving frame is aligned with the accumulation tool and will be different from the original one.Figure 6: An illustration of the coordinate systems and main parameters of the prototyping system.y o position of the moving frame. The second part, q, is the quaternion of the moving frame after the rotation from its initial position. Both of them are defined in terms of the coordinate system of the base frame. Accordingly we know:z,y,x o)q.

Consequently the displacement(cid:74)(cid:74)(cid:74)(cid:74)(cid:74)(cid:71) PO O BBPy= (x o),z,;position of the tool tip;is the position of the ball joint in the movingooPPPilΔ of the ith strut is:Δ = l i(cid:74)(cid:74)(cid:74)(cid:74)(cid:71) PB i iB−l i0where 0il is the initial joint offset.In our hardware design, the base joints are not exactly located on the legs. Consequently the deviation of the lead screw from the base joint needs to be considered when computing the displacement. That is,Δ = l i(cid:74)(cid:74)(cid:74)(cid:74)(cid:71) PB i iB2−dev2−l i0(4)4.2 Movement SimulationA simulation software system was developed for computing the tool path based on Equations [1]- [4]. Figure 7.a shows the Graphical User Interface (GUI) of the developed simulation software system. A G-code file with defined quaternion at each point can be opened in the view area.

Figure 7.a shows the Graphical User Interface (GUI) of the developed simulation software system. A G-code file with defined quaternion at each point can be opened in the view area. The position of each joint can be set in the left panel. The parameters values are set based on the machine calibration, which will be discussed in Section 5. Before the simulation, an axis motion command file will be generated from the G-code file to define quaternion based on Equations (3) and (4). The simulation starts when the axis motion command file passes the constraint check. After that, the final file will be sent to the motioncontrol software system, whose GUI is shown in Figure 7.b.

The simulation starts when the axis motion command file passes the constraint check. After that, the final file will be sent to the motioncontrol software system, whose GUI is shown in Figure 7.b. Accordingly, the 6 linear motors will be controlled with the defined moving positions and speeds.(a) The simulation software system (b) The motion control software system Figure 7: Software systems developed for the parallel kinematic machine.4 Platform Calibration(cid:74)(cid:74)(cid:74)(cid:74)(cid:71) and 0il (i=1~6) need to be set in the coordinate P iO P transformation models as described in Equations (3) ~ (4). The calibration of these parameter values is necessary for more accurate control of the 6-axis motion. The calibration approach that is used in our system is to iteratively update the system parameters by an error model such that a defined cost function can be minimized based on the given orientation and translation of the moving platform.

The calibration approach is presented in more details as follows.(cid:74)(cid:74)(cid:74)(cid:74)(cid:74)(cid:71) B iO BBPA set of 42 parameters including,4.1 Cost Function and Error ModelA cost function for the 6-axis platform calibration is defined in [19]. It uses strut length measurement residual at all measured poses as the objective function.

Since 42 system parameters are unknown in the axis displacement generation model,the measurement of seven or more poses are required. In our experiments, we randomly chose twelve poses, twice the minimum number of poses, to ensure the accuracy of the calibrated parameter values.A computer vision based method is used in measuring the tool tip position and orientation. As shown in Figure 8, two cameras are positioned orthogonal to each other. Three targets were placed on the moving platform. Target 1 is located at the tool tip. Targets 2 and 3 were symmetric about the X axis of the moving frame. Each camera is calibrated such that the 3D point positions of the targets can be computed based on the captured 2D images of both cameras. Accordingly, the rotation matrix of the moving frame can be derived by the SVD method based on the positions of the three targets.

Accordingly, the rotation matrix of the moving frame can be derived by the SVD method based on the positions of the three targets. The process is described in more details as follows.Figure 8: An illustration of the platform pose measurement using two cameras.4.2.1 Camera CalibrationThe two cameras are first calibrated in order to identify the relation between a pixel on a 2D image and its corresponding 3D position. Considering the nonlinearity of a camera, we equally divide the working volume into 11 layers along the X and Y axes for camera 1 and 2, respectively. The distance between two neighboring layers is 5mm. A printed chessboard is placed at each layer; an image will be taken to analyze all the corner points of the chessboard. Since the coordinates of each corner point on the chessboard are known, a database of the relations between an image pixel and its base frame coordinate at each layer can be established.

Since the coordinates of each corner point on the chessboard are known, a database of the relations between an image pixel and its base frame coordinate at each layer can be established. For example, suppose the ith corner point on the chessboard at layer j in ) in the camera 1’s calibration volume has world coordinate (1 x y z ). After its pixel ( ij1 ij1 ij,,x image 1 1 1 x y z , , ij ij ij,y image,x imageimage is identified, a database for camera 1 can be obtained in the form of (,y imageInversely, if pixel position of a point on the image is given as (x image,y image), its corresponding 3Dposition at layer j can be calculated through the bilinear interpolation of the four corners of the checker box in which the pixel falls.4.2.2 3D Point Coordinate ComputationFor a certain platform pose, the three targets on the moving platform can be captured by the two )cameras (refer to Figure 9).

Accordingly, two groups of pixel positions ( 1 x image1 y image) and (x,2image,y2imagefrom cameras 1 and 2 can be recorded for any target. Since the depth of the target in the camera’s view volume is unknown, we can calculate its 3D position at each layer first. For example, the world coordinates P1,0 can be retrieved from the database related to pixel ( ) at layer 0. Similarly all these 3D coordinates identified at each layer will form a distorted camera view line (refer to Figure 9).1 x image1 y image,).For the same target, the two camera lines may not exactly intersect due to the camera calibration errors. Instead, the 3D coordinate of the target can be computed as the point with the minimum distances to both camera lines. For example, the closest distance between the two camera lines as shown in Figure 9 is between segment P1,1P1,2 and segment P2,5P2,6.

For example, the closest distance between the two camera lines as shown in Figure 9 is between segment P1,1P1,2 and segment P2,5P2,6. Such an approach was also used in [20].Figure 9: Point retrieval by calculating the closing point between two camera lines.5 Multi-directional AM Process OverviewBased on the developed low cost parallel kinematic machine, the major steps of a multi-directional AM system include (1) tool path generation, (2) system coordinate transformation, (3) platform constraint checking, (4) movement simulation, and (5) part fabrication. In the tool path generation, an input CAD model is sliced in desired building orientation. The slicing direction can be determined based on surface normal or feature skeleton. A sliced layer contour can be sampled into a set of discrete points. Each sampling point can be used to compute the positions of the six linear actuators. At each point, the tool position and orientation are given.

Each sampling point can be used to compute the positions of the six linear actuators. At each point, the tool position and orientation are given. As discussed in Section 3, the 3D coordinate of the tool along with its accumulation orientation is converted into a transformation model to compute the related distances L0-L5 for the six struts of the machine. The displacement vector can then be sent to the motion controller to achieve the desired platform pose.Note that not all the vectors computed based on the transformation process can be achieved by the prototype system. Several constraints exist including the limitation of actuators’ stroke (i..e [Lmin, Lmax]), the limitation of the range of the passive joint, and the minimum distance between each actuator’s lead screw. Hence the computed results need to be checked using the known constraints.

Hence the computed results need to be checked using the known constraints. The tool path can be simulated in our software system to verify whether the 6-dimensional motion vectors satisfy all the constraint equations. Finally, the verified 6 dimension coordinate vectors (L0, L1, L2, L3, L4, L5) can be sent to the control system to start the building process.Figure 10 shows the data processing pipeline of the developed multi-directional AM system. A CNC control software system (e.g. Mach 3) can be used in converting a given CAD model into numeric control G-codes. The generated G-codes have only the coordinate positions along which the tool tip will travel. The tool can be positioned at various orientations for a given position in the G-code.

The generated G-codes have only the coordinate positions along which the tool tip will travel. The tool can be positioned at various orientations for a given position in the G-code. One good candidate is to align the tool orientation with the related part surface normal.Multi-Direction SlicingTransformationSamplingCAD ModelTool Path PlanningParallel Kinematic Machine Motion CommandL0, L1, L2, L3, L4, L5Building Part by AM (FDM or CNC Accumulation)SimulationConstraint CheckFigure 10: Data processing pipeline of the multi-directional AM systems using the parallel kinematic machine.Figure 11: The prototype system for the multi-directional FDM process.6 Experimental TestsThe developed parallel kinematic machine can be used in various multi-directional AM processes such as the FDM or CNC accumulation processes. A photo of the developed prototype system with a filament extruder of the FDM process is shown in Figure 11.

A photo of the developed prototype system with a filament extruder of the FDM process is shown in Figure 11. In this section, some of the test cases we performed are presented to demonstrate its capability.In the first test case, a CAD model as shown in Figure 12.a is to be fabricated. The planned tool paths are shown in Figure 12.b. Note that the surface normal of the shape changes from the vertical direction n1 to n2. For the horizontal portion of the model, n1 is chosen as the building direction. For the slope portion of the model, n3 is chosen as the building direction for better attachment to the previously deposited layers. Figure 13 shows the building results by using the multi-directional FDM system. The melted filament can attach well to the former solidified layers. Hence the slope portion can be built without any supports by using the material deposition direction of n3 instead of n1.(a)(b)Figure 12: A model with an inclined plane to be built.

Hence the slope portion can be built without any supports by using the material deposition direction of n3 instead of n1.(a)(b)Figure 12: A model with an inclined plane to be built. (a) Different building directions; (b) planned tool path.(a)(b)Figure 13: Test results. (a) Part fabrication by FDM extrusion in a tilted angle; (b) built part.In the second test case, three characters “USC” are to be built on a tilted plane as shown in Figure 14.a. In the linear motion based FDM machines, a nozzle can only extrude filaments vertically. Hence it would be difficult to build features on a tilted plane since a large gap will exist between the tilted plane and the vertical nozzle in order to avoid the collision between them. In comparison, a desired gap distance can be achieved by rotating the nozzle to be perpendicular to the plane. Accordingly the characters can be fabricated on the tilted plane, which are shown in Figure 14.b.Figure 14: Test results.

Accordingly the characters can be fabricated on the tilted plane, which are shown in Figure 14.b.Figure 14: Test results. (a) Problem with the vertical building on a tilted plane; (b) built characters on the plane.The third case is to deposit a curved line on the surface of a bottle. As shown in Figure 15, the bottle body is a cylinder with a diameter of ~31mm. In order to build a curved line along its surface, the nozzle direction needs to be constantly changed in order to keep the tool to have a building direction that is coaxial with the surface normal. The nozzle orientations at the start and end positions are shown Figure 15.a. The angles of the nozzle will change from -20º to +20º during the building process. Figure 15.b shows the fabricated result.(a)(b)At startAt endFigure 15: Test results.

The angles of the nozzle will change from -20º to +20º during the building process. Figure 15.b shows the fabricated result.(a)(b)At startAt endFigure 15: Test results. (a) FDM nozzle at different positions; (b) built result on the curved surface.A youtube video of the developed multi-directional AM system and the related experimental tests canbe found at [21].7 Conclusion and Future WorkIn the paper, a low-cost 6-axis motion system based on the Stewart mechanism has been developed for multi-directional additive manufacturing. Our design is modular and can easily be incorporated with various accumulative tools. In addition, the system based on six linear actuators is relatively inexpensive. It can move tools in satisfactory speed. By enabling full 6-axis motions between a tool and a work piece, an AM system can deposit materials from multiple directions.

It can move tools in satisfactory speed. By enabling full 6-axis motions between a tool and a work piece, an AM system can deposit materials from multiple directions. We further demonstrated that the drawbacks associated with the layer-based AM processes such as the limitations on building-around- inserts can be overcome by the multi-directional AM processes.The specifications of our 1st generation multi-direction AM machine are:(1) Working volume. The working volume of our system is designed as 300mm×300mm×150mm. The working volume depends on the travel range of the lead screws and slide bushings, as well as the angle range of the ball joints.(2) Accuracy. The accuracy of the prototype system is <0.5mm when the tool direction remains the same, e.g. building planar features on a tilted plane; however, the error can be larger when the tool direction keeps changing, e.g. building features on a curved surface.

building planar features on a tilted plane; however, the error can be larger when the tool direction keeps changing, e.g. building features on a curved surface. A main reason is the poor calibration resolution we used (11 planes were used with a layer distance of 5mm). Another possible improvement is to improve motion planning algorithm and the hardware construction. (3) Rotation angle of the platform. The rotation angle of the moving frame in our prototype system is ±30º. Like the working volume, the rotation range is mainly dependent on the selected ball joints, and the ravel ranges of lead screws and slide bushings.We plan to further improve the accuracy of the prototype system by better software and hardware implementation, and more precise calibration. In addition, test cases for applications such as part repairing and building around inserts will be further studied.AcknowledgementWe acknowledge the help of Tom Kerekes from Dynomotion Inc.

In addition, test cases for applications such as part repairing and building around inserts will be further studied.AcknowledgementWe acknowledge the help of Tom Kerekes from Dynomotion Inc. for sharing his 3-axis parallelkinematic machine design and the use of KFlop control board.Reference[1] Y. Pan, X. Zhao, C. Zhou, Y. Chen. Smooth Surface Fabrication in the Mask Projection basedStereolithography. SME Journal of Manufacturing Processes, Vol. 14, No. 4, pp. 460-470, 2012.[2] Sager B. SLA characterization for surface finish improvement: Inverse design methods for process planning[dissertation]. 2006.[3] Benay Sager, David W. Rosen. Use of parameter estimation for stereolithography surface finish improvement.Rapid Prototyping Journal. Vol.14, No. 4, 213-20, 2008.[4] Pandey PM, Venkata Reddy N, Dhande SG. Improvement of surface finish by staircase machining in fuseddeposition modeling. Journal of Materials Processing Tech. Vol. 132(1):323-31, 2003.[5] Mason A.

velopment of a Melt Pool Tracking Vision System for Laser DepositionTodd Sparks1,a,Lie Tang1,b, Frank Liou1,c 1Department of Mechanical and Aerospace Engineering Missouri University of Science & Technology 400 W. 13th St., Rolla, MO 65401, USA atsparks@mst.edu, bltx8d@mst.edu, cliou@mst.eduAbstractThis paper chronicles the development of a vision system for tracking melt pool mor- phology in the laser metal deposition process. This development is to augment an existing temperature feedback control system. Monitoring both the temperature and shape of the melt pool is necessary because of the eﬀects of local geometry on the cooling rate at the melt pool. Temperature feedback alone cannot accommodate this eﬀect without complex process planning. The vision system’s hardware, software, and integration into the laser deposition system’s controller is detailed in this paper.

The vision system’s hardware, software, and integration into the laser deposition system’s controller is detailed in this paper. Preliminary testing and the eﬀects on deposition quality is also discussed.1Introduction and MotivationThere are numerous structural components for air and vehicle frames, propulsion systems, and other industrial applications that are made from oversized forgings or from castings that have high raw material content. These components are complex in shape and can be made from expensive alloys. Current manufacturing processes for these components are costly and labor-intensive, exhibit long manufacturing lead times, and require complex tooling. The maturation of rapid manufacturing method will reduce lead time for part production and improve remanufacturing capabilities. Therefore, there is a strong need to directly produce and repair complicated functional parts to a known and repeatable condition.

Therefore, there is a strong need to directly produce and repair complicated functional parts to a known and repeatable condition. The following trends have already helped bring interest in the concept of rapid manufacturing:Smaller lot sizes / More segmented product oﬀerings (mass customization) • Shorter development cycles • Shorter program/platform life • Rising material costs • Environmental pressures (governmental and social) • Desire for part consolidation (DFM/DFA)1Despite these trends, no process currently exists which satisﬁes industrial needs in this changing marketplace. Only a few niche applications currently use rapid manufacturing techniques for production parts. This is possible due to two factors:Tradition - The current industrial work force is trained and familiar with the traditional sub-tractive manufacturing methods.Quality - Additive processes do not have the history behind them that the traditional subtrac- tive processes do.

Repeatability of both geometry and material characteristics for rapid manufacturing processes are still not well understood.This work attempts to address an issue associated with quality concerns. Consistent volumetric addition via laser deposition requires some knowledge of the current state of the melt pool. Real- time tracking of melt pool temperature is impossible through contact measurement and diﬃcult via non-contact methods. This work presents an alternative approach which attempts to use the melt pool size rather than temperature to control the process.2 BackgroundSince its appearance, rapid prototyping technology has been of interest to various industries that are looking for a process to produce a part directly from a CAD model in a short time. Among them, direct metal deposition process is one of the few process which directly manufactures a fully dense metal part without intermediate steps.

Among them, direct metal deposition process is one of the few process which directly manufactures a fully dense metal part without intermediate steps. In this process, metallic powder is injected into a laser generated heat spot where the material is melted and forms a melt pool which quickly solidiﬁes into metal layers [3, 4, 9]. Parts are built to completion layer by layer, from bottom to top. This process is similar to other rapid prototyping technologies in its approach to fabricate a solid component by layer additive methods.Funded by the National Science Foundation and Air Force Research Laboratory, the Missouri University of Science & Technology (MS&T) has developed the Laser Aided Manufacturing Process (LAMP) [5–8,10–14]. The system, illustrated in ﬁgure 1, combines both deposition and machining in a single setup. This eliminates part re-setup which is a signiﬁcant advantage over processes which require post-processing in separate machinery.

This eliminates part re-setup which is a signiﬁcant advantage over processes which require post-processing in separate machinery. Any production part will have a speciﬁcation for both its geometric and mechanical properties. Microstructural characteristics are a major factor in determining the performance of a material. Solidiﬁcation microstructure is a function of the temperature gradient in both space and time [2, 15]. The part geometry determines the heat conduction properties of the structure as it is being built, so it has an eﬀect on the thermal history of the part, and thus the solidiﬁcation microstructure. Constant conditions are necessary for consistent processing. Consistent conditions require some kind of feedback, thus the impetus for this work.3Imaging Design ConceptsThe primary challenge for image capture in this project is the reﬂection of the laser beam oﬀ of the shiny surface of the melt pool. Three successive concepts were tested for dealing with this issue.

Three successive concepts were tested for dealing with this issue. The ﬁrst two are attempts at selecting what portion of the spectrum of light coming from2Figure 1: The LAMP process during a repair operation - The laser cladding head is shown on the right. A touch probe head is currently held in the machining spindle. A repaired H13 tool steel die component is held in the vise.the melt pool reach the CCD sensor. The third design looks at a single wavelength of light that is well outside the expected response of the melt pool. The sensor used for this study is a 1.3 megapixel CCD from the commonly available Microsoft VX1000 webcam.3.1 Filtering and Selective ReﬂectionThe ﬁrst imaging concept tested was a simple ﬁltering design. The ﬁlter stack consists of a piece of heat proof glass to remove IR and a short pass ﬁlter to cut oﬀ wavelengths longer than 700nm.

The ﬁlter stack consists of a piece of heat proof glass to remove IR and a short pass ﬁlter to cut oﬀ wavelengths longer than 700nm. A 300µm pinhole lens was used rather than the webcam’s included optics as a way to further cut down on the signal reaching the CCD. Figure 2(a) shows the conﬁguration of this design.The second concept, illustrated in Figure 2(b), uses a more aggressive approach to remove the laser reﬂection from the image. A piece of heat proof glass cuts out some the IR. Then a cold mirror reﬂects visible light through a green dichroic ﬁlter and a polarized glass disc.Both the ﬁrst and the second designs resulted in image acquisition such as the image shown in Figure 3. Measurement techniques for the captured images are discussed below in Section 4. The melt pool is concealed by the bright bloom in the center of the picture. The second bright spot in the upper right of the image is the reﬂection of the melt pool on the copper nozzle tip.

The melt pool is concealed by the bright bloom in the center of the picture. The second bright spot in the upper right of the image is the reﬂection of the melt pool on the copper nozzle tip. This is obviously not a good solution for imaging the melt pool, so a third option was devised.3.2 IlluminationTo avoid the imaging problems associated with the previous two designs, the third imaging concept seeks to eliminate all of the thermal radiation from the melt pool. According to black body radiation theory, as illustrated in Figure 4, the melt pool should never radiate light in the UV range. Thus, that is where this design concentrates.To image the melt pool using light in the UV range, a narrow bandpass ﬁlter was used to cut the camera’s range down to 389nm to 399nm. Since this is outside the range of visible light as well the melt pool’s thermal response, an external illumination is needed.

Since this is outside the range of visible light as well the melt pool’s thermal response, an external illumination is needed. Three UV (390nm - 395nm)3(a) Pinhole lens holder and ﬁlter attachment for the simple ﬁltering concept.(b) Assembly of the reﬂective design showing the heat resistant glass (clear), the cold mirrior (dark blue), and the polarized glass (gray).Figure 2: Solidworks 2008 assemblies of the ﬁrst two concepts.LEDs were used. Figure 5 shows the conﬁguration of the Illumination design. Unfortunately, this design has not yet been tested as of the time of this writing.4Image Processing and System IntegrationImage processing for this project is done using the Python bindings for Intel’s open source computer vision library, OpenCV.

This library provides the functionality to both capture images through the VX1000’s video4linux driver interface and do the necessary operations to locate and measure the melt pool within the image.Capturing a clear image of the melt pool during laser deposition can be diﬃcult due to the powder obscuring the image. Others have attempted to solve this issue with a fuzzy thresholding algorithm [1]. However, if the input is a video stream, simply taking the median value of each pixel from a set of frames will eﬀectively erase fast moving objects from the image. This is similar to a trick used in image processing to erase tourist from a photo by using this operation on a set of stills taken over the span of a few minutes.To measure the melt pool, a best ﬁt ellipse is ﬁt to a binary image, with the major and minor axes of the ellipse being the information of interest. The orientation of the camera is then used to do a perspective correction.

The orientation of the camera is then used to do a perspective correction. Ideally, multiple cameras will capture simultaneous images from diﬀerent angles to achieve a good approximation of the melt pool shape. This information is then passed to the LAMP system’s National Instruments Real-Time controller via a serial USB connection.4Figure 3: This is a sample result from the imaging system. The blue ellipse is a best ﬁt ellipse over what it thinks is the melt pool. The red and green lines are the major and minor axes, respectively.Figure 4: The CIE 1931 XYZ Color Space relates black body radiation color and temperature. (Image courtesy of Wikimedia Commons)5 Discussion and ConclusionDuring this work, all of the structural components for the concept imaging designs were fabricated from ABS plastic on a Dimension FDM machine.

When testing these designs, it became apparent that the white ABS plastic did very little to block the intense near-IR and IR radiation coming from the melt pool. In fact, the image was completely saturated even when the melt pool was not in the camera’s ﬁeld of view. This problem was remedied by shielding the camera with Aluminum. A system for measuring melt pool geometry is a necessary component for a robust laser de- position process. This work represents steps taken to implement such a system. Unfortunately, equipment problems in LAMP lab at MS&T prevented the completion of this project in time for inclusion into this paper.5Figure 5: Conﬁguration of the UV Illumination design6 AcknowledgmentsThis research was supported by the National Science Foundation grants DMI-9871185 and IIP- 0637796, and a grant from the U.S. Air Force Research Laboratory contract # FA8650-04-C- 5704.

Air Force Research Laboratory contract # FA8650-04-C- 5704. The support from Boeing Phantom Works, Product Innovation and Engineering, LLC, Spartan Light Metal Products Inc, MS&T Intelligent Systems Center, and MS&T Manufacturing Engineering Program, is also greatly appreciated.References[1] Mathew Asselin, Ehsan Toyserkani, Mehrdad Iravani-Tabrizipour, and Amir Khajepour. De- velopment of trinocular ccd-based optical detector for real-time monitoring of laser cladding. In Proceedings of the IEEE International Conference on Mechatronics & Automation, July 2005.[2] William Hofmeister, Melissa Wert, John Smugeresky, Joel A. Philliber, Michelle Griﬃth, and Mark Ensz. Investigation of solidiﬁcation in the laser engineered net shaping (lensTM) process. Technical report, Sandia National Labs, 1999. Sandia National Labs report.[3] J.L. Koch and J. Mazumder. Rapid prototyping by laser cladding.

Liou. Automatic process planning of a multi-axis hybrid manufacturing system. In DETC, September 29 - October 2 2002.[11] Jianzhong Ruan and Frank W. Liou. Automatic toolpath generation for multi-axis surface In Proceedings of the 2003 ASME Designmachining in a hybrid manufacturing systemg. Automation Conference, Chicago, Illinois, September 2-6 2003. Paper No. DAC-48780.[12] Jianzhong Ruan, Jun Zhang, and Frank W. Liou. Support structures extraction for hybridlayered manufacturing. In DETC, 2001.[13] Todd Sparks, Vinay Kadekar, Gail Richards, Frank W. Liou, Venkat Allada, Ming Leu, Faisal Anam, and Siddharth Shinde. An advanced manufacturing workshop for high-school teachers and students. In Proceedings of the 2005 ASEE Annual Conference & Exposition, Portland, Oregon, June 12-15 2005.[14] Todd Sparks, Vinay Kadekar, Yogesh Thakar, Frank W. Liou, and Ashok K. Agarwal. Edu- cating high school students and teachers in rapid prototyping and manufacturing technologies.

Liou, and Ashok K. Agarwal. Edu- cating high school students and teachers in rapid prototyping and manufacturing technologies. In Proceedings of the 2004 ASEE Annual Conference & Exposition, June 20-23 2004.[15] Jos´e E. Spinelli, Ot´avio Fernandes Lima Rocha, and Amauri Garcia. The inﬂuence of melt convection on dendritic spacing of downward unsteady-state directionally solidiﬁed sn-pb al- loys. Materials Research, 9(1):51–57, 2006.7.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDevelopment of a Method to Derive Design Guidelines for Production-suitable Support Structures in Metal Laser Powder Bed Fusion S. Lammers*,†, T. Lieneke*,†, D. Zimmer*,†Chair of Design and Drive Technology, Paderborn University, 33098 Paderborn, Germany †Direct Manufacturing Research Center (DMRC), Paderborn University, 33098 Paderborn, GermanyAbstractSolid support structures in metal laser powder bed fusion have a significant influence on the economic applicability, component quality and process stability and represent a central challenge for widespread industrial use. As the connection of the components to the building platform by supports is essential, the negative effects must be minimized at the same time as the supporting effect is optimized.

As the connection of the components to the building platform by supports is essential, the negative effects must be minimized at the same time as the supporting effect is optimized. Within the scope of this study, a standardized method is developed that allows the investigation of several support structures and parameters with regard to their influence on the target variables: component quality, process efficiency and stability. In ad dition to the proof of general suitability, the applicability is investigated using so-called standard elements. Based on the experimental results, design guidelines are derived, which will serve as a basis for decision-making during the selection of support structure for an individual application.Introductionto conventional, subtractive manufacturing processes, additive manufacturing involves building up components element by element or layer by layer, without tools. [1].

[1]. The layer-by- layer build up is based on the underlying three-dimensional CAD model, which is converted into a sequence of two-dimensional layer models [2].In contrastThe simplification of the three-dimensional problem into many two-dimensional layers induces design freedom. In this way, structures can be realized that would be difficult or impossible to produce using conventional manufacturing processes [3]. In a ddition to complex external free-f orm surfaces, internal structures such as cooling channels close to the contour, structure-o ptimizing grid fillings or force-flow-optimized sandwich structures can also be utilized using additive manufacturing processes [4]. The manufacturing of topology-optimized or bionic structures is also possible [5, 6]. These advantages result in a large lightweight design potential and consequently a large potential for material- efficient production as well as resource-saving operation of additively manufactured components.

In addition, function integration can reduce the number of components required to fulfill different functions and reduce the assembly effort, which enables cost reduction and further weight reduction. [7, 8, 9, 10]. The Metal Laser Powder Bed Fusion (LPBF) process, among others, is established for the additive manufacturing of metallic components.However, LPBF implicates process-specific restrictions and disadvantages. For example, the process is characterized by a high proportion of manual, time-intensive work steps. In addition to the support structure generation and data preparation, which take place before production, these also include downstream processes such as cooling, separation from the building platform, powder removal and support structure removal.

The support structure-related steps are particularly inefficient, but unavoidable, because support structures are needed in the LPBF process for a robust build process, to minimize warpage, to fix the components during the process, and to dissipate the process-induced heat. After the build process, however, these support structures are largely unwanted and must be removed at great expense. Since they also increase material requirements and production times, they are considered as a necessary evil. Other post- processing methods such as heat treatment, machining or surface finishing may also be necessary to improve dimensional accuracies, surface qualities, porosity and mechanical properties inherent to the process. Compared to conventional processes, the low production speed for large series is also a limiting factor. In addition, the component size is limited by the available build chamber of the ad ditive manufacturing system [3; 11].

In addition, the component size is limited by the available build chamber of the ad ditive manufacturing system [3; 11]. It is unknown, how long these224disadvantages will remain. However, the highly dynamic market in the field of additive manufacturing and extensive research projects give reason to hope for significant improvements [12, 13, 14, 15].MotivationIn order to minimize or completely eliminate the inherent disadvantages of the LPBF process andthe support structures, profound fundamental knowledge is required. To this end, design guidelines for users have been developed in recent years, which provide support for the production suitable design of components for additive manufacturing [4, 12, 16, 17, 18] or assistance for the finishing suitable design of additive components [13, 15, 19]. Design guidelines with respect to optimized support structure generation are also partially available [14, 20, 21, 22].

Design guidelines with respect to optimized support structure generation are also partially available [14, 20, 21, 22]. However, these often only deal with the variation of a support structure parameters and its influence on the component properties or on the process. This paper therefore focuses on the development of a standardized method for the quality assessment of support structures with respect to different component and process properties under variation of support structure parameters. The method should enable the user to determine the influence of different support structure parameters, such as hatching, tooth geometry or fragmentation, on the component or the process. As a result, it should be possible to derive design guidelines for successful support structure design and to benchmark different support structures.

As a result, it should be possible to derive design guidelines for successful support structure design and to benchmark different support structures. These design guidelines with reference to support structures are to be called production-suitable support structure design guidelines in the following.Support structure parametersThe most common support structure in use is the block support shown in Figure 1. It is a support structure composed of narrow walls of wall thickness bW in a rectangular grid pattern. The block support is used to support voluminous and massive components as well as for smaller components [14]. The center-to-center spacing aH of adjacent support walls in the rectangular grid is called the hatch distance.Figure 1: Relevant block support parameters. At the top the side view with perforated side of contour walls.

At the top the side view with perforated side of contour walls. At the bottom the top view on the support structure fragmentation.225The contact area to the component and the build platform is usually designed with so-called teeth. The tooth geometry is described by the support structure parameters top length lT and tooth height hZ. Fragmentation is another support structure parameter. It divides the lattice scheme in the xy-plane into adaptable segments. Perforations, on the other hand, introduce interruptions in the support structure walls in the z-direction. The objective in the support structure parameter selection is to design a support structure that has sufficient resistance to residual stresses and process forces, while allowing better peelability and powder removability.

The parameters hatch distance aH, wall thickness bW, top length lT and the tooth height hZ mentioned here have the greatest influence on these properties and will be used for the further investigations [8, 23].Method to derive design guidelines for production-suitable support structuresThe standardized procedure of the method to derive design guidelines for production-suitable support structures is described in the following. It is intended to enable a comparison of different variants of support structures with regard to criteria of component quality, process efficiency and s tability. In addition, the usefulness in practical applications, such as inclined overhangs or internal contours, will be examined more in detail.

In addition, the usefulness in practical applications, such as inclined overhangs or internal contours, will be examined more in detail. The method is divided into the five steps shown in Figure 2.Figure 2: Schematic structure of the method to derive design guidelines for production-suitable support structures.The first step is to identify relevant quality criteria for support structures that influence component quality, process efficiency and stability. These criteria must be determined for different support structures in order to be able to perform a quality assessment. In a second step, test specimens will be developed which can be supported by different support structures and allow the investigation of relevant quality criteria. In order to keep the test effort low, small, material-saving test specimens were developed for pre-probing, which allow an efficient investigation of a large number of different support structures in a few build jobs.

In order to evaluate geometries that are closer to the application, further test specimens were developed. These are based on standard elements already used in ADAM [16]. These so-called application-oriented test specimens are limited to support structures for whom a basic suitability could be demonstrated during the pre-probing. The test specimens are dimensioned in such a way that they can be manufactured on common equipment with a building platform size of at least 250 x 250 mm. In the third step, the two test specimen types are manufactured226using the LPBF process. In the fourth step, the test specimens are examined with regard to the quality criteria for support structures identified in step 1. The test results are entered in a clear graphical evaluation, which shows the effect and interaction of different support structure parameters on the quality criteria.

The test results are entered in a clear graphical evaluation, which shows the effect and interaction of different support structure parameters on the quality criteria. In the final step, design guidelines are derived based on the results to provide designers with basial rules for application- adapted support structure design. The method is structured to be able to integrate further boundary conditions such as production systems, materials, production parameter sets or support structure parameters.Step 1: Identification of support structure quality criteriathe documented experience of the AM community in the form of scientific publications and field reports. They are marked in a bold font in this chapter.The quality criteria were selected onthe basis of previousinvestigations andAs a basic quality criterion, a visual inspection of the imaging accuracy is first performed. The actual geometry is compared with the target geometry.

The actual geometry is compared with the target geometry. Any conspicuous features, such as fractures, powder deposits or deformations, are recorded.In the investigation of the attainable component quality, the measurement of the shape and dimensional accuracy of the test specimen is carried out in all three spatial directions. For the dimensional stability in the z-direction, t he maximum distortion at the bottom of the test specimen is determined. If the deformation in the positive z-direction in the area of the uppermost layer exceeds the value of the layer thickness during the process, the process can be aborted due to the collision of the recoater with component areas [22]. It is known that the risk of collision of the recoater comes from the deformation at the current top of the component. However, the strongest deformation in the z-direction is present at the bottom side of the components.

However, the strongest deformation in the z-direction is present at the bottom side of the components. If this deformation is subcritical, the shape deviation is usually compensated by the further coating and exposure process and thus no longer occurs on the top side of the finished component. For this reason, the measurement of dimensional accuracy in the z-direction is limited to the bottom side of the test specimen. By determining the dimensional accuracy in the z-direction, a possible sagging of the lowest test specimen layer in the further course of the process is detected. Preliminary investigations have shown that successive shrinkage in the z-direction can be observed in some support structures.

Preliminary investigations have shown that successive shrinkage in the z-direction can be observed in some support structures. This results in a positive dimensional deviation in the z height of the component and a negative dimensional deviation at the support structures, as the contact area between the support structure and the test specimen shifts in the negative z direction, while compensation takes place on the upper side due to the continuous coating. Another reason for a dimensional deviation in the z-direction can be a sinking of the melt into the powder bed or the undesired sintering of powder particles on the underside of the component [8].When measuring the shape and dimensional accuracy in the xy-plane, deviations in the test specimen width and depth are to be investigated. As the layers cool, the material contracts, causing tensile residual stresses in the top layer and induced compressive residual stresses in the layers below.

As the layers cool, the material contracts, causing tensile residual stresses in the top layer and induced compressive residual stresses in the layers below. This can induce dimensional deviations of the component width and worsen or prevent the connection to the support structures. Furthermore, the machining of the components to revise the dimensional deviations reduces the process efficiency [22].In addition to shape and dimensional accuracy, surface roughness influences component quality. The measurement of surface roughness is limited to the underside of the test specimen, since this represents the contact surface to the support structure. Negative effects on the surface quality are to be expecte d from remaining support residues as well as from adhering and not completely melted pow der particles [21].

Negative effects on the surface quality are to be expecte d from remaining support residues as well as from adhering and not completely melted pow der particles [21]. In addition to the negative influence on the component quality, a pronounced surface roughness often necessitates further post processing, especially in the area of fits and functional surfaces, which reduces the process efficiency [4, 8, 9].When designing support structures, a compromise must be found between sufficient strength and the removability of support structures [7]. For the industrial use of LPBF, the process efficiency must be increased by reducing the manual post-processing effort. In this context, simplified removability of the support structures is of particular importance [3, 23].In the case of software-supported automatic support structure generation (e.g.

In this context, simplified removability of the support structures is of particular importance [3, 23].In the case of software-supported automatic support structure generation (e.g. via Materialise Magics), in which adjustments of individual parameters may have to be made, the data preparation time varies227depending on the number of necessary settings. If the support structures are even created completely manually or a unit cell is designed for cellular supports, a significantly greater time requirement can be assumed. In addition to the data preparation time, the necessary data volume is also recorded. For example, a massive increase in storage capacity can be expected for very complex, filigree support structures with large surfaces. This poses challenges for the data infrastructure.

For example, a massive increase in storage capacity can be expected for very complex, filigree support structures with large surfaces. This poses challenges for the data infrastructure. In addition, delays in the process chain are to be expected due to an increased time requirement for slicing the build job and loading it on the LPBF system.Material consumption also determines the efficiency of the process. This is often opposed to the requirement for sufficient mechanical resistance of the support structures [22]. Both the material consumption of the support structures and the amount of powder which are enclosed by the support structures and cannot be removed via spaces between the teeth, perforations or fragmentations must be taken into account.

This residual powder poses a risk of contamination and may not be reusable due to contamination or oxidation [5].Step 2: Development of test specimens for the investigation of the quality criteria under consideration of the test effortTwo different test specimen geometries are used to pre-test a large number of different parameter combinations for support structures. These are referred to as PK1.1 and PK1.2 in the following. The basic support structure quality criteria defined in step 1 can be investigated with these. A small number of test specimens with reduced material consumption is desirable without reducing the significance or reproducibility of the results.PK1.1 is a simple cuboid basic geometry, which is shown in Figure 3. The simple test specimen geometry is intended to minimize any influence of the test specimen geometry on the effects of the support structures.

The simple test specimen geometry is intended to minimize any influence of the test specimen geometry on the effects of the support structures. Furthermore, the small test specimen dimensions are suitable for manufacturing as many test specimens as possible in one build job. Thus, 81 parameter combinations can be investigated with one build job on widely used machines with a minimum build platform size of 250 x 250 mm, which corresponds to a full factorial test plan with four factors of three factor levels each. This experiment can be easily and meaningfully extended. In addition, the expected distortion usually starts from the corners, since these are particularly exposed areas, as the surrounding powder makes heat dissipation difficult [24].

In addition, the expected distortion usually starts from the corners, since these are particularly exposed areas, as the surrounding powder makes heat dissipation difficult [24]. Overall, despite its small dimensions, the square base is a geometry that is particularly susceptible to distortion and on which meaningful results can be expected for the various combinations of support structure parameters.Below the test specimen, the examined support is inserted with a height of 4 mm, which corresponds to a widely used minimum support structure height.

For the measurements on the test specimens, a clamping in form of an 8 mm x 8 mm cuboid was manufactured, the influence of this cuboid on the quality criteria can be precluded because of its subsequent production during the build job.After the production of PK1.1, a measurement of the shape and dimensional accuracy in the xy-plane and the positional accuracy of the test specimen top in the z-direction on the building board is carried out. This is followed by an assessment of the removability of the support structures. After the support structure has been removed, the test specimens are examined for the dimensional accuracy of the bottom, since the highest distortion can be assumed on there. In addition, the dimensional accuracy in z-direction can be determined. Subsequently, the surface roughness of the bottom of the test specimen is measured in order to investigate the influence of the support structure residues and the fracture surfaces.

Subsequently, the surface roughness of the bottom of the test specimen is measured in order to investigate the influence of the support structure residues and the fracture surfaces. Furthermore, the time required for data preparation and the resulting data volume can be determined for the test specimens PK1.1.The PK2.1 test specimen, also shown in Figure 3, is used to investigate material consumption. Its basic geometry was used by GRALOW ET AL. for corresponding tests and has proven to be suitable [20]. Due to its tetrahedral base plate, PK2.1 has a predetermined breaking point in the contact area with the building platform. This prevents a serious deviation of the total weight due to a varying test specimen volume after separation from the building platform. On the square base plate with an edge length of 12 mm, supports are inserted which allow a connection to the final cover plate.

On the square base plate with an edge length of 12 mm, supports are inserted which allow a connection to the final cover plate. Support structures with different parameter combinations are inserted in the area between the cover plate and the base plate. The cover plate thus simulates the component to be supported and the base plate the building platform. The area filled with support structures has a different removability of the residual powder depending on the used support structure parameters and is thus suitable for determining the material consumption considering different tooth geometries, perforations or228fragmentations.

To assess the suitability of the most promising support structures, various test specimens are developed, each representing a specific application situation. Standard elements are used, which have already been established in other scientific work [13, 15, 16]. For this purpose, non-curved, single-curved and double-curved standard elements with different attribute characteristics are investigated and the effect of support on these widely used geometric elements is investigated (PK3.1 to 3.5). In addition, removability under restricted accessibility is evaluated (PK3.7) and the change in the top length of the teeth as a function of the downskin angle of the supported surface is examined more in detail (PK3.6).The PK3.1 test specimen shown in Figure 4 was developed to evaluate the support effect that can be achieved on externally supported, non-curved base elements. It consists of differently inclined overhangs in two different length and width steps.

It consists of differently inclined overhangs in two different length and width steps. Non-curved overhangs are application-oriented geometries, which simulate downskin surfaces having the same normal vector at every point. Thus, the influence of the downskin angle on different quality criteria will be investigated.Figure 4: Test specimen PK3.1 - external non-curvedThe angle of the overhangs referring to the building platform is 0°, 20° and 40° and thus covers the range between a horizontal orientation of the downfacing surface of the component and the self-supporting angle, above which the use of solid support structures is unnecessary. Different literature values exist for this purpose. According to CALIGNANO, component surfaces must be supported below a relative angle of 30°, GRALOW ET AL. state a value of 40° and according to ADAM the limit value is 45°. For the following test specimens, an angle of 40° is chosen as the maximum value of the orientation [16, 20, 23].

state a value of 40° and according to ADAM the limit value is 45°. For the following test specimens, an angle of 40° is chosen as the maximum value of the orientation [16, 20, 23]. In addition to the orientation of the overhangs, their width and length are varied. With increasing length, the path for dissipation229of thermal energy increases [16]. Increased negative influences of gravity and residual stresses can be detected at overhangs during the process [21]. With test specimen PK3.1, overhangs with a length of 10 mm and 30 mm projected onto the building platform are investigated. For the support of the overhangs, their width is also decisive. This is varied between 5 mm and 20 mm. The solid center bar ensures robust fabrication and subsequent clampability in the test fixture. After production, a measurement of the shape and dimensional accuracy is carried out on the build platform.

After production, a measurement of the shape and dimensional accuracy is carried out on the build platform. The scientific investigations are carried out using the Nikon Altera® 8.7.6 coordinate measuring machine with the Renishaw® PH10M Plus rotary swivel head, the Renishaw® SP25M probe holder and the Renishaw® SP25_L40_D2 probe. After removing the support structures from the build platform, the measurement is repeated. In this way, any further distortion due to residual stress can be determined. In addition, a Keyence VR3200® macroscope is used to check the shape and dimensional accuracy of the cross-sectional profile of the overhanging beams and to examine the surface roughness of the downskin surfaces in accordance with DIN EN ISO 4288 and DIN EN ISO 3274 [16, 26].Internal non-curved contours are also investigated in order to simulate application cases such as cavities or cooling channels [4].

For this reason, the influence of the support structures on differently inclined, internal, non-curved base elements is investigated on the test specimens PK3.2.1, PK3.2.2 and PK3.2.3 shown in Figure 5.Figure 5: Test specimen PK3.2.1 with 0° orientation, PK3.2.2 with 20° orientation and PK3.2.3 with 40° orientationThe considered orientations correspond to 0°, 20° and 40° for the non-curved standard elements on external surfaces. The projected length of the surface to be supported is also identical at 10 mm and 30 mm, respectively. Spacings of 10 mm and 20 mm between the vertical walls were selected for the internal support structures. The wall thickness is 2 mm. The removability of the support structures from the internal, non- curved contours is investigated on the test specimens.

The wall thickness is 2 mm. The removability of the support structures from the internal, non- curved contours is investigated on the test specimens. After the support structure has been removed, a measurement of the cross-sectional profile is made and the surface roughness of the supported surfaces is examined.In addition to the non-curved standard elements, many components also show single-curved standardelements in their geometry. Analogous to the test specimen PK3.1 for the examination of non-curved, external standard elements and PK3.2 for the examination of non-curved internal standard elements,230examinations of single-curved standard elements are carried out on the test specimens PK3.3 and PK3.4.1, PK3.4.2 and PK3.4.3. Here, the same quality criteria are examined using the same measuring techniques.

Here, the same quality criteria are examined using the same measuring techniques. The Specimens are shown in Figure 6.Figure 6: Test specimen PK3.3 - external single-curved (top) and test specimen PK3.4.1 to 3.4.3 - internal single- curved (bottom).Other widely used standard elements are double-curved standard elements. The PK3.5 specimen shown in Figure 7 is used to investigate the support quality on convex and concave geometries. Both the inward-curved (concave) and outward-curved (convex) double-curved standard elements are examined with the three different diameter grades of 5 mm, 10 mm, and 20 mm.

Both the inward-curved (concave) and outward-curved (convex) double-curved standard elements are examined with the three different diameter grades of 5 mm, 10 mm, and 20 mm. Height images are taken on the double- curved standard elements placed on specimen PK3.5 for the qualitative evaluation of the shape and dimensional accuracy as well as for the assessment of remaining support structure residues.Figure 7: Test specimen PK3.5 – external and internal double curved.The investigation of the test specimen PK3.6 shown in Figure 8 is an extension of the investigationson test specimen PK3.1. The contact situation between the component and the support structure is usually described by the component geometry. This changes the angle of intersection with the support structures and causes a change in tooth geometry. For this purpose, the PK3.6 specimen has a downskin surface with four subdivisions, which have a downskin angle of 10°, 20°, 30° and 40°. The width of the specimen is 5 mm.

For this purpose, the PK3.6 specimen has a downskin surface with four subdivisions, which have a downskin angle of 10°, 20°, 30° and 40°. The width of the specimen is 5 mm. Here, after separation from the build platform, an optical examination of the top length for the different downskin angles can be carried out with the aid of a macroscope.231Figure 8: Test specimen PK3.6 – downskin angle.The PK3.7 specimen shown in Figure 9 is used to investigate the support removability with one-sided and two-sided accessibility. It is divided into two sections. On the left side, there is accessibility to the internal support structure on two sides, while on the right side the support is surrounded by solid walls on five sides. The test specimen is first removed from the building platform and then examined regarding removability. For this purpose, the time required to remove the support structure with a least possible residue is determined.

For this purpose, the time required to remove the support structure with a least possible residue is determined. As qualitative evaluation criteria, the necessary use of tools is determined, and the remaining support and powder residues are examined.Figure 9: Test specimen PK 3.7- support removability.Step 3: Production of the test specimens under variation of the supportPK1.1 and PK1.2 are manufactured with 81 different parameter combinations of the support structures. A separate build job is provided for each of the test specimens PK1.1 and PK2.1. The positioning of the test specimens with the different parameter level combinations for build job 1 is shown in Figure 10 on the left. The orientation of the test specimens is parallel to the recoater lip. This orientation is indispensable for the subsequent metrological examination on a coordinate measuring machine and is justifiable in terms of production technology due to the small component dimensions.

In addition, the orientation ensures maximum loading of the test specimens by the recoater, thus putting the durability of the support structures against process forces to the test [4]. The distance between the test specimens is 10 mm to facilitate measurement and detachment and to avoid mutual interferences between the support structures and test specimens due to the arrangement. The build jobs can be manufactured on machines with a build platform size of 250 x 250 mm or larger. A reference block is inserted for measurement by means of a coordinate measuring machine. The build jobs are produced in multiple realizations, considering randomization and block formation to eliminate global changes in boundary conditions during production, such as ambient temperature or powder charge, as well as local thermal differences across the build platform from the results. Figure 10 on the right shows the position names on the build platform.

Figure 10 on the right shows the position names on the build platform. Between several realizations, the assignment of the different parameter level combinations to the position names has to be randomized. The fabrication was carried out on a SLM Solutions SLM 250HL system using the material 316L (1.4404) and the parameter set 316L_SLM_MBP3.0_ 50_CE2_400W_Stripes_V2.0.232Figure 10: Realized arrangement of the test specimens parameter for build job 1(left) and general positionames (right).PK3.1 to PK3.6 are also manufactured under the described boundary conditions. However, not all 81 paremeter combinations of the support structures are considered here. This would increase the test effort and the costs of the test specimen production too much. Instead, the best parameter combination from the pre- probing at PK1.1 and PK2.1 are selected and used to support PK3.1 to 3.7.

Instead, the best parameter combination from the pre- probing at PK1.1 and PK2.1 are selected and used to support PK3.1 to 3.7. In this context, PK3 represents test specimen geometries close to the application, which allow validation of the results from the pre-probing and were developed based on established standard elements [16]. The arrangement of PK 3.1 to PK3.7 is shown in Figure 11. In addition, a problem-specific adaptation of the build job can be performed by removing test specimens not related to the application to reduce the experimental effort. Further on, an area for freely selectable test specimens is provided.

Further on, an area for freely selectable test specimens is provided. This allows to place test specimens for the calibration and verification of a numerical simulation or tensile tests.Figure 11: Arrangement of the test specimens PK3.233Step 4: Examination and evaluation of the test specimens with regard to the quality criteriaAs an example, the examination of shape and dimensional accuracy in the xy -plane is listed below and the graphical evaluation is described.

The investigations for the other quality criteria were evaluated analogously, so that diagrams as shown in Figure 12 exist for each quality criterion.Figure 12: Overall result plot of shape and dimensional accuracy in the xy plane (maximum dimension below deviation).To investigate the shape and dimensional deviation in the xy plane, eight measuring points were probed on the PK1.1 specimen with the Nikon Altera® 8.7.6 coordinate measuring machine, two on each of the four side faces with an offset of 1 mm in the z direction from the top and bottom edges of the specimen. For evaluation, the maximum dimensions below the nominal dimension was calculated. Figure 12 shows the maximum dimension below for all investigated parameter level combinations. Here, the abscissa shows the hatch distance of the support structures in three parameter levels and the ordinate shows the wall thickness of the support structures.

Here, the abscissa shows the hatch distance of the support structures in three parameter levels and the ordinate shows the wall thickness of the support structures. Each smaller, colored square in the coordinate system represents a separate additional diagram in which the tooth geometry was varied on the basis of tooth height and top length. In this overall result display, all possible combinations of hatch distance, wall thickness, top length and tooth height can be quickly compared with each other. The quality criterion maximum dimension below deviation in the xy- direction is quantified by the color scheme. The dark green areas thus represent parameter level combinations that exhibit a particularly low dimensional deviation. Reddish shades represent a greater dimensional deviation.

The dark green areas thus represent parameter level combinations that exhibit a particularly low dimensional deviation. Reddish shades represent a greater dimensional deviation. Detailed values can be found in the legend.The overall result plot in Figure 12 shows a clear correlation between the wall thickness parameter and the maximum dimension below of the specimen width in the xy plane. Whereas with a wall thickness of 0.25234mm or 0.4 mm the dimension below does not exceed a value of -0.160 mm, a wall thickness of 0.2 mm (depending on the choice of the other parameters) shows a deviation of up to -0.424 mm. This variant of the block support at this location has a hatching of 2 mm, a top length of 0.4 mm and a tooth height of 0.7 mm. Overall, there is an increased dimension below with reduction in wall thickness. This can be explained by a decrease in deformation resistance, which cannot prevent contraction of the solidifying and cooling material.

This can be explained by a decrease in deformation resistance, which cannot prevent contraction of the solidifying and cooling material. When comparing the measurement results in the bottom row of the overall results plot, an increase in the dimension below with an increase in the size of the hatching also becomes apparent. This can be seen in the comparison of the three support variants, which have a wall thickness of 0.2 mm, a top length of 0.4 mm and a tooth height of 1.3 mm. No comparable influence can be seen with a wall thickness of 0.25 mm or 0.4 mm. Obviously, the deformation resistance of the single support wall at corresponding wall thicknesses is sufficient to compensate the lower number of support walls when the hatch distance is increased. Based on the bottom row, the influence of the top length can also be observed.

Based on the bottom row, the influence of the top length can also be observed. It can be clearly seen that with an increase in the contact length of the teeth to the supported surface, a reduction in the maximum dimension below can be achieved. This effect can also be attributed to an increase in deformation resistance due to the larger total contact area. No comparably clear correlation is shown for the tooth height. While an increase in the dimension below with an increase in the tooth height can be observed for a wall thickness of 0.2 mm and a hatching of 2 mm (right field of the bottom row), this influence cannot be seen for a center distance of 1 mm or 1.5 mm. With a greater wall thickness of 0.25 mm or 0.4 mm, there is also no influence of this parameter. An increase in tooth height appears to have a negative effect only with a simultaneous weakening of the support structure due to a reduced wall thickness and an increased hatch distance.

An increase in tooth height appears to have a negative effect only with a simultaneous weakening of the support structure due to a reduced wall thickness and an increased hatch distance. The effects and interactions of the individual support structure parameters on the quality criterion can be seen in the effect diagram in Figure 13.Figure 13: Effect diagram for different parameters for dimensional accuracy in the xy plane (maximum dimension below deviation).A comparison of the four investigated support structure parameters shows that wall thickness (B) has most significant influence on the quality criterion dimensional accuracy in the xy-plane. If the wall thickness is reduced from the medium level of 0.25 mm to the minimum possible value of 0.2 mm, the maximum dimensional deviation increases by -0.157 mm on average. If, in the opposite direction, the parameter is 235increased to a value of 0.4 mm, there is a decrease in the dimensional undershoot of 0.022 mm.

If, in the opposite direction, the parameter is 235increased to a value of 0.4 mm, there is a decrease in the dimensional undershoot of 0.022 mm. The influence of the hatching (A) and the tooth height (D) is comparatively small in the considered value ranges of the parameters with 0.038 mm and 0.012 mm, respectively. In both cases, an increase in the dimension below can be observed when the parameter is increased. The top length (C) of the teeth shows a more significant influence. Increasing this parameter from 0.4 mm initially to 1 mm reduces the dimension below in the xy plane by 0.080 mm on average. However, comparing all parameters, increasing the wall thickness can be recommended as the most effective measure to reduce the dimension below. The dominance of effect B by adjusting the wall thickness can also be seen from the 2-factor interactions.

The dominance of effect B by adjusting the wall thickness can also be seen from the 2-factor interactions. The course of the 2-factor interaction AB, BC and BD corresponds approximately to that of effect B.Step 5: Derivation of design guidelinesDesign guidelines can be derived on the basis of the results determined in step 4. In the following the design guidelines for the quality criterion of dimensional accuracy in the xy-plane are exemplarily presented in Table 1. Design guidelines for other quality criteria were derived analogously. For this purpose, the results of the effect calculations for the individual support structure parameters are summarized in written form and visualized with the aid of a drawing “suitable for production” and one “not suitable for production”. To determine the listed limit values aHmax, bWmin, lTmin and hZmax, use the overall result representation in Figure 12.

The limit values are therefore dependent on the selected support structures and on the boundary conditions such as machine, material and production parameters.Table 1: Design guidelines for production-suitable support structures with regard to the quality criteria shape and dimensional accuracy in the xy plane (maximum dimension below deviation).Hatch distance of support structures (Effect A)Description: The hatch distance aH of adjacent retaining walls should be chosen as small as possible to minimize the dimension below in the xy-plane.Suitable designUnsuitable designWall thinkness of support structures (Effect B)Description: The wall thickness bW of support walls should be chosen as to as large minimize the dimension below in the xy-plane.possibleSuitable designUnsuitable design236Top length of support teeth (Effect C)Description: The top length the support teeth should be chosen as large as possible to minimize the dimension below in the xy planelT ofSuitable designUnsuitable designHeight of support teeth (Effect D)The Description: height hZ of the support teeth should be chosen as low as possible to minimize the dimension below in the xy planeSuitable designUnsuitable designAdditional influences through effect interactionsThe wall thickness (parameter effect B) dominates over the other parameter effects The increase of the wall thickness represents the most efficient measure to reduce the dimensional undershoot in the xy-planeConclusion and outlookOne of the hurdles to more widespread industrial use of the LPBF process is the unavoidable use of solid support structures.

Their utilization, caused by a variety of tasks they fulfil, affects part quality and process efficiency. To counteract the disadvantages and to gain a comprehensive process understanding with regard to support structures, this publication deals with the development of a method to derive production- oriented design guidelines for support structures. Based on application-relevant quality criteria such as shape- and dimensional accuracy, surface roughness, removability, data preparation time, data volume and material consumption, test specimens were developed and manufactured. To examine the test specimens, a standardized procedure was developed to show the influence of various support structure parameters on the quality criteria. As a result, the designer receives a graphical representation of the influence that each parameter has onto the quality criteria.

As a result, the designer receives a graphical representation of the influence that each parameter has onto the quality criteria. This graphical representation can be used as a well-founded basis for the selection of an application-specific suitable support structure as well as for the deriving of design guidelines.To extend the results, it is possible to investigate further support structures or to vary further boundary conditions such as the manufacturing equipment, the material or the manufacturing parameters. Due to the modular structure of the method, this is feasible in an easy way. For example, the procedure can be used to build up a database for support structures.

velopment of a Multi-material Mask-Image-Projection-based Stereolithography for the Fabrication of Digital MaterialsChi Zhou, Yong Chen*, Zhigang Yang, Behrokh Khoshnevis Epstein Department of Industrial and Systems Engineering University of Southern California, Los Angeles, CA 90089Corresponding author: yongchen@usc.edu, (213) 740-7829Abstract Digital materials such as the ones shown by Objet’s Connex family demonstrate that a new material with desired characteristics can be achieved by combining two different base materials with various concentrations and structures. We investigate the feasibility of using additive manufacturing processes based on digital mask projection in the fabrication of such digital materials. A multi-material mask- image-projection-based Stereolithography process has been developed. The related challenges on the development of such a process are identified. Our approaches to address such challenges are presented.

The related challenges on the development of such a process are identified. Our approaches to address such challenges are presented. A testbed has been developed to fabricate component with designed digital materials. Experimental results illustrate desired material properties can be achieved based on the developed process. Several multi- material designs have been produced to highlight the capability of this promising technology for fabricating three-dimensional, multi-material objects with spatial control over placement of both material and structure. The limitations and challenges for future development have also been identified.Keywords: Additive manufacturing, multi-material fabrication, bottom-up projection, part separation.1. IntroductionLayer-based additive manufacturing (AM) is a collection of techniques for manufacturing solid objects by the sequential delivery of energy and/or material to specified points in space to produce that solid.

Differentiated from conventional manufacturing processes, a unique capability of the AM processes is that multiple materials or functionally gradient materials can be added in a single component during the building process. An example of such multi-material AM systems is the OBJET Connex series (www.objet.com). Based on its PolyJet Matrix Technology, these three-dimensional (3D) printers are capable of manufacturing digital materials with complex internal structures. By combining two base materials in specific concentrations and structures, as many as 51 different materials can be created in a single printed part [1]. Hence product components can have material designs with desired mechanical properties, e.g. both soft and hard materials can be embedded in products such as tooth brushes and remote controllers.

Hence product components can have material designs with desired mechanical properties, e.g. both soft and hard materials can be embedded in products such as tooth brushes and remote controllers. Such fabrication capability also opens up exciting new options that were impossible before.Our work is motivated by the recent 3D printer development especially by the digital material fabrication. Note the OBJET Connex machines are based on jetting model materials from designated micro-scale inkjet printing nozzles. Such a process has inherent limitations on the selection of base materials since the jetted liquid need to have certain properties on viscosity and curing temperatures in order to be jetted. To significantly expand the material selections, we investigated the feasibility of other AM processes for the digital material fabrication.

To significantly expand the material selections, we investigated the feasibility of other AM processes for the digital material fabrication. Besides the inkjet printing technology, the fused deposition modeling (FDM) process can naturally be extended for fabricating parts out of multi-materials since FDM has separate extrusion nozzles for the build and support materials. Khalil et al. [2] presented a multi-nozzle deposition system for producing 3D tissue-engineered scaffolds. However, the FDM process has limitations on its minimum nozzle size and is relatively slow. Hence it is not suitable for digital material fabrication. There have also been attempts at using selective laser sintering (SLS) for multi-material fabrication [3-7]. However, accurate material feeding and recoating required by the digital material fabrication is difficult to be designed in the SLS process.In this research we focused on another representative AM process, Stereolithography Apparatus (SLA).

By using a laser and liquid photocurable resin, SLA has been shown to be able to offer high quality surface finish, dimensional accuracy, and a variety of material options. To address its speed limitation, our research focused on the mask-image-projection-based Stereolithography (MIP-SL) process instead. An illustration of the MIP-SL process is shown in Figure 1. Different from a laser used in SLA, a Digital Micromirror Device (DMD) is used in the MIP-SL process to dynamically define mask images that will be projected on a surface area. A DMD is a microelectromechanical system (MEMS) device that enables one to simultaneously control ~1 million small mirrors to turn on or off a pixel at over 5 KHz. In the MIP-SL process, the three-dimensional (3D) CAD model of an object is firstly sliced by a set of horizontal planes. Each thin slice is then converted into a two-dimensional (2D) mask image. The planned mask image is then sent to the DMD.

Each thin slice is then converted into a two-dimensional (2D) mask image. The planned mask image is then sent to the DMD. Accordingly the image is projected onto a resin surface such that liquid photocurable resin can be selectively cured to form the layer of the object. By repeating the process, 3D objects can be formed on a layer-by-layer basis. Compared to the laser-based SLA, the MIP-SL process can be much faster due to its capability of simultaneously forming the shape of a whole layer. Two test parts built by our prototype MIP-SL system by using two different materials are also shown in the figure.Figure 1. An illustration of the MIP-SL process.2. Principles of a Mask-Image-Projection-based SL System for Digital Material FabricationMultiple vats are required for different types of liquid resin in the multi-material SLA and MIP-SL processes. As a natural extension to the single material SLA system, Maruo et al.

As a natural extension to the single material SLA system, Maruo et al. [8] first presented a multiple material stereolithogarphy system by manually removing the vat from the platform and draining the current material, rinsing the vat, returning the vat to the platform, and dispensing a prescribed volume of a different material into the vat. However, based on the lengthy manual material changing process, the system was limited to simple 2.5D microstructures. Wicker et al. [9-11] extended the work by developing a multiple vat carousel system to automate the building process including washing, curing and drying cycle between build materials. Based on similar ideas, Choi et al [12] reported a multi-material MIP-SL system for fabricating micro-scale objects. Arcaute et al. [13] and Han et al.

Based on similar ideas, Choi et al [12] reported a multi-material MIP-SL system for fabricating micro-scale objects. Arcaute et al. [13] and Han et al. [14] also presented an automatic material switching approach by dispensing the solution using a pipette into a custom-made small vat, and subsequently washing the current solution before changing to the next solution. Based on the technique, fabricated 3D scaffolds for heterogeneous tissue engineering have been demonstrated.A core challenge in the use of multiple materials in SL is how to manage material contamination between changing different materials used in the fabrication process. The previous research [8-14] on developing multi-material SLA and MIP-SL systems are all based on the top-down projection. As shown in Figure 2, in order to accommodate the part size in Z direction, a large tank has to be maintained for keeping the resin level.

As shown in Figure 2, in order to accommodate the part size in Z direction, a large tank has to be maintained for keeping the resin level. Due to the deep vat, draining and cleaning the current resin before changing to another resin vat takes long time and leads to significant material waste. To address the problem, Kim et al. [15, 16] presented a process planning approach to minimize the material changeover number for a given multi-material CAD model. That is, if different materials are separated in a CAD model, one material can be fully or as much as possibly built before transferring to another material. Even being able to reduce the material changeover efforts, the approach, however, is not general especially for digital material fabrication, in which different materials are interlocked with each other.Figure 2.

An illustration of the multi-material SL process based on top-down projection.Functional microstructures especially digital material fabrication requires the development of a general MIP-SL process similar to the polyjet process that can fabricate all combinations of multiple resins. The main challenge to be addressed in such a multi-material MIP-SL system is to reduce the material waste and increase cleaning efficiency during the resin tank switching process. To address the problem, we investigated the bottom-up projection in the multi-material MIP-SL process. An illustration of such a system is shown in Figure 3. The light source is projected from the bottom of the transparent vat. Since the current build layer is formed at the bottom of the platform, the container depth is independent on the part height. Thus the liquid in the vats can potentially be as shallow as a layer thickness.

Thus the liquid in the vats can potentially be as shallow as a layer thickness. When switching resin tank, only the portion of the built model that contact liquid resin need to be cleaned. Thus the material changeover efforts can be significantly reduced with less material waste.Figure 3. An illustration of the multi-material SL process based on bottom-up projection.Our efforts on developing such a bottom-up projection based MIP-SL system are presented for digital material fabrication. The remainder of the paper is organized as follows. Part separation study is presented in Section 3. Our approach to reduce part separation force is presented in Section 4. Transition between multiple tanks is presented in Section 5. The experimental setup for performing physical experiments is discussed in Section 6. The experimental results of multiple test cases are presented in Section 7. Finally conclusion with future work is drawn in Section 8.3.

The experimental results of multiple test cases are presented in Section 7. Finally conclusion with future work is drawn in Section 8.3. Part Separation Study of the Bottom-up Projection based MIP-SL ProcessIn the bottom-up projection based MIP-SL process, a cured layer is sandwiched between the previous layer and the resin vat. The solidified material may adhere strongly to the corresponding rigid or semi- rigid transparent solidification substrate, causing the object to break or deform when building platform moves away from the vat during the building process. One approach to conquer the attaching force is to increase the exposure to significantly over-cure current layer such that its bonding force with previous layers can be increased. However, over curing, in the same time, brings the problem of bad surface quality and inaccurate dimensions. Another approach to address the problem is to apply certain coating on the resin vat such that the attaching force can be reduced.

Another approach to address the problem is to apply certain coating on the resin vat such that the attaching force can be reduced. Suitable coatings include Teflon and silicone films can help the separation of the part from the vat [17, 18]. A coated Teflon glass has also been used in the machines of Denken [19] and EnvisionTec [20].Even with the intermediate material, the separation force can still be rather large. Huang and Jiang [18] investigated the attaching force for the coating of an elastic silicone film. Based on a developed on- line force monitoring system, test results indicate that the pulling force will increase linearly with the working area. The experiments indicate that, for a square of 60×60mm, the pulling force to separate the part from the film is bigger than 60 N.

The experiments indicate that, for a square of 60×60mm, the pulling force to separate the part from the film is bigger than 60 N. Such a big attaching force between the cured layer and the vat is a key challenge to be addressed in the development of the bottom-up projection based MIP-SL system.In our research, another type of coating material, Polydimethylsiloxane (PDMS, Sylgard 184, Dow Corning), is applied on the resin vat. This is based on a unique property of the PDMS film during the polymerization process that was identified in [21]. Dendukuri et al [21] presented a photolithography- based microfluidic technique for continuously fabricating polymeric particles. The developed technique is based on the oxygen-aided inhibition near the PDMS surfaces to form chain-terminating peroxide radicals. Hence a very thin oxygen inhibition layer (~2.5 μm) will be formed that can prevent cured layer attach to the PDMS film. We studied the part separation forces based on the PDMS film.

Hence a very thin oxygen inhibition layer (~2.5 μm) will be formed that can prevent cured layer attach to the PDMS film. We studied the part separation forces based on the PDMS film. The experimental results are discussed as follows.Figure 4. Part separation study in the MIP-SL process.3.1 Separation Forces for Solidified Resin based on the PDMS FilmA set of physical experiments have been performed to investigate the separation force of a cured layer based on a coated PDMS glass. The setup for measuring the pulling force is shown in Figure 4.a. Two FlexiForce sensors (Tekscan, South Boston, MA) with a range of 0-25 lbs are sandwiched between the fixture and vat. Since the vat is free at the bottom and the side, and only fixed at the top, the pulling forceby the part will be transferred to the sensors when the platform levels up. The two sensors are connected to a microcontroller, which can sample and record the sensors’ readout in over 50Hz.

The two sensors are connected to a microcontroller, which can sample and record the sensors’ readout in over 50Hz. In the experiments, we first use a given mask image to build a certain number of layers (e.g. 25 layers). We then begin to record the separation force in the building process of the next few layers. For each layer, after the designed mask image has been exposed for a certain time, the platform will be raised up slowly and the related readouts of the sensors are then recorded.The three factors that could affect the separation force and were considered in our study include (1) exposure time; (2) image area; and (3) image shape. Designed experiments based on these factors are tested to understand their effects. Figure 4.b shows a set of mask images that have been used in the experiments for testing the effect of image shapes. The tested projection patterns, including triangle, square, hexagon, circle, t-shape, u-shape, band, and star-shape, have the same area in each test.

Pulling forces of cured layer from a PDMS film in different test cases.Based on the experiment results, it can be observed that: (1) As the Z stage moves up, the separation force will increase until it reaches a peak value whenthe cured layer is detached from the PDMS film;(2) The peak force gets bigger when the same mask image is exposed longer; (3) The peak force gets bigger when a larger image area is projected; (4) The image shape has some effects on the peak force but not as significant as the exposure timeand the projection area;(5) With the coated PDMS film on the vat, the separation force is still considerably large (~100 ozor 27.8 N for an image area of 625mm2 with 1 second exposure).3.2 Separation Force for Liquid Resin without CuringA similar set of experiments are conducted to analyze the pulling force of a part without liquid resin being cured between the part and the vat.

In the experiments, we first use an image of a square (A = 625 mm2) to build a certain number of layers (e.g. 25 layers). Based on it, the built part is lowered to form a certain gap with the PDMS film. Without exposing any image to cure liquid resin, the platform is then raised up slowly and the related separation forces are recorded form the force sensors. Different gap values (0.1- 0.5mm) have been tested. The experiment result is shown in Figure 6. It can be seen that:(1) The separation force is smaller than the related cases with solidified resin; (2) The separation force decreases with a bigger gap size between the part and the PDMS film; (3) The separation force can only be neglected until the gap size is larger than 0.5 mm.Figure 6. Pulling forces of a built part from a PDMS film for different gap sizes.4.

Pulling forces of a built part from a PDMS film for different gap sizes.4. Two Channels Design for the Bottom-up Projection based MIP-SL ProcessThe large separation force between the cured layer and the resin vat may fail the building process when the bonding force between the current layer and previous layers is smaller than the separation force. In addition, the PDMS film will have cracks after building multiple layers due to the cyclic loading and related material fatigue.To facilitate the bottom-up projection based MIP-SL process, a new two-channel design has been developed, which can fundamentally address the large separation force in the building process. The developed approach is motivated by the following two observations:(1) As discussed in [18], the intermediate of cure layer and silicone film becomes a vacuum stage when the resin is sufficiently cured.

Position 2). Note the vacuum has been broken at this moment. Hence, if the PDMS film is thick enough (e.g. >0.5mm), the part can now be easily separated from the vat. After moving up the part by a certain distance d (position 3), the vat is moved back such that the part is on top of the channel with PDMS (position 4). Finally the platform will move down by a distance d – layer_thickness for the building of a new layer. Note the motion in the X axis can be added on the vat and related frame. Hence the accuracy of the MIP-SL system is not affected by the X translation since there is no relative motion between the platform and the projection device.An appropriate thickness of the coated PDMS film can be determined by considering the followingfactors:(1) The thickness of oxygen inhibition layer (around 2.5μm) on the PDMS surface is independent on the thickness of the PDMS film [21].

Thus the shear force is not directly related to the PDMS film thickness.(2) More light energy will be consumed by the film if the PDMS film is thicker; (3) The PDMS film should be thick enough such that the gap between the cured layer and the vatsurface at Position (2) will be big enough to have a small separation force.(4) More resin has to be maintained for a thicker PDMS film in the channel without PDMS. Considering all the above factors, the PDMS film thickness is set at 1mm in our prototype system.Based on it, the separation and shearing forces of the two channel design are discussed as follows.4.1 Separation Forces for Solidified ResinTo verify the proposed two-channel system design, physical experiments that are similar to the ones in Section 3.1 have been conducted. The same setup as shown in Figure 4.a has been used in measuring the separation forces. The same mask images as shown in Figure 4.b have been used in building test layers.

The same setup as shown in Figure 4.a has been used in measuring the separation forces. The same mask images as shown in Figure 4.b have been used in building test layers. By repeating the same experiments for the two-channel system design, the testing results are shown in Figure 8. In each figure, the curves record both the sliding and pulling-up stages. As can be seen from the figures, the force in Z direction is very small during the sliding of the resin vat. During the platform pulling-up stage, the peak separation forces are also relatively small (around 2-4 oz or ~0.83 N). Themeasured forces are only 4-5% of the related ones observed in the one channel design.

Themeasured forces are only 4-5% of the related ones observed in the one channel design. In addition, the variations of the exposure time, the image area and the image shape have a smaller effect on the separation forces in the new design.(a) T = 1 sec, Area = 625 mm2(b) T = 0.5 sec, Area = 625 mm2Sampling time (×20 msec)Sampling time (×20 msec)(c) T = 0.15 sec, Area = 625 mm2(d) T = 1 sec, Area = 625/4 = 156 mm2Figure 8. Pulling forces of cured layer for two channels system in different test cases.4.2 Shearing Forces in the X Axis The FlexiForce sensors were used in a modified setup to measure the shearing force in the X axis. However, no meaningful readouts were recorded form the sensors due to the small shearing force. To quantitatively estimate the value of the shearing force, a set of square rods with different sizes were built using the two channels system. The built rods are shown in Figure 9. Their heights are 10mm and the minimal rod size is 0.3×0.3mm.

The built rods are shown in Figure 9. Their heights are 10mm and the minimal rod size is 0.3×0.3mm. Note we also successfully built rods with even smaller sizes. However, the rods were so fragile that they lost the mechanical strength to sustain themselves when the part was taken out from the resin vat and washed in isopropyl alcohol.Nevertheless, for a rod with a size of 0.3×0.3mm, the upper bound on the tangential force that can be added on it can be analytically estimated. As shown in Figure 9, the testing rods in the experiment can be modeled as a cantilever beam. Suppose the length of the beam is L, the size of the beam section is b*b, the force in tangent direction is F. The maximum bending stress occurs at the end can be calculated as:σ =/Mc I, where I is the section modulus,Ib=4 /12, andcb=/ 2. Put them together, we haveσ =6FL 3 b.Suppose the allowable blending stress is []σ and the minimal beam section size is [ ]b.

Put them together, we haveσ =6FL 3 b.Suppose the allowable blending stress is []σ and the minimal beam section size is [ ]b . We will have thefollowing equation:F≤3 [ ][ ] b σ L 6. The parameters for this experiment are listed as follows. []σ =65MPa,[ ]b =0.3mm, L=10mm. According to the equation, the upper bound of the tangential force is only 0.03N or 0.11oz. Compared with the separation force in Z direction, the shearing force in X direction is rather small.Figure 9. Shearing force verification5. Transition between Multiple TanksAs discussed before, the challenge of using multiple materials in the MIP-SL process is managing the contamination between different materials. The proposed two-channel system enables a small separation force in the bottom-up projection. Hence shallow vats can be used in the MIP-SL process to reduce the material waste and the required cleaning effort.

Hence shallow vats can be used in the MIP-SL process to reduce the material waste and the required cleaning effort. To ensure no contamination between different resin vats, different cleaning strategies have also been explored and identified.5.1 Shallow Vat StudyIt is desired to have as little liquid as possible in a resin vat to reduce the contact of the part and liquid resin. However, when the thickness of liquid resin in a tank is too small, islands that have no liquid will appear on the bottom surface due to surface tension. Hence the minimum thickness of liquid resin on the PDMS surface needs to be determined based on the tested resins. A scaled syringe was used to gradually inject resin into the two-channel tank until the resin can fully cover the whole PDMS surface. As shown in Figure 10, the related thickness for Perfactory SI500 (yellow color) resin is found to be ~0.5mm.

As shown in Figure 10, the related thickness for Perfactory SI500 (yellow color) resin is found to be ~0.5mm. Reducing the viscosity of resin can reduce the surface tension and accordingly the minimum resin thickness. During the building process, a pump can be used to dynamically add liquid resin into the vat to compensate for the material consumption.Figure 10. Minimum thickness of resin on the PDMS surface.5.2 Cleaning Resin Residue on Built LayersLiquid resin may accumulate around the perimeter of the object and the bottom of the cured layer when it is raised up from the vat. To avoid material contamination when changing to different resin vats, the excessive materials on the bottom and the side of part surface should be removed before building a new layer. Various cleaning approaches have been tested.

Various cleaning approaches have been tested. The best candidate we identified is a two-stage cleaning strategy based on:(1) Rough cleaning: a soft brush is moved relative to the part, which can remove the majority of liquid resin on the bottom and the perimeter of the part. The resin collected in the brush tank can be recycled to refill the building tank.(2) Final cleaning: Due to the surface tension, resin residue can still be found on part surface after the rough cleaning. To thoroughly clean the resin residue, ultrasound cleaning was identified as the most effective approach for final cleaning. After immersing the bottom portion of the built part in a container with liquid solvent (e.g. 90% isopropyl alcohol and 10% water by volume), high frequency ultrasound can be provided. Due to the applied ultrasound, microscopic bubbles in the solvent will be formed on the part surface and imploded under the pressure of agitation.

Due to the applied ultrasound, microscopic bubbles in the solvent will be formed on the part surface and imploded under the pressure of agitation. The generated shock waves will impinge on the part surface. Consequently the resin can be quickly and thoroughly rinsed in all directions. The approach is especially effective for resin inside small cavities, which is difficult to be removed by other cleaning methods.After the final cleaning, the part is wetted with solvent and must be dried up before merging into another material. Otherwise, a new layer can not be bounded well with the previous layers. In our prototype system, a fan is used to blow dry air on the part to dry out the alcohol residual.

Otherwise, a new layer can not be bounded well with the previous layers. In our prototype system, a fan is used to blow dry air on the part to dry out the alcohol residual. After the part is dry, the building process may resume and layers with a new type of material can be added.For two types of materials (A and B), the required stations in our prototype system consisting of two resin vats, two brush tanks, a ultrasound cleaner, and a fan are shown in Figure 11.a. Even though the shallow vat enables only a small amount of material being cleaned in the system, the cleaning procedure will take the majority of the cycle time, which significantly reduces the throughput of the whole process. Hence reducing the material switchover number is still important in the multi-material MIP-SL system.rDriershAnCBrush BFigure 11. Hardware system of the prototype multi-material MIP-SL system.6.

Hardware system of the prototype multi-material MIP-SL system.6. Experimental Setup6.1 Hardware SystemA prototype system has been built for verifying the presented methods. The hardware setup of the developed multi-material MIP-SL system is shown in Figure 11.b. In the designed system, an off-the- shelf projector (CASIO XJ-S36) was used as the projection device. The use of a commercial projector cansignificantly reduce the prototype cost and simplify the system design. The optical lenses of the projector were modified to reduce the projection distance. Various projection settings including focus, key stone rectification, brightness and contrast were adjusted to achieve a sharp projection image on the designed projection plane. The DMD resolution in our system is 1024×768 and the envelope size is set at 48×36 mm. A linear stage from VELMEX Inc (Bloomfield, NY) is used as the elevator for driving the platform in the Z axis.

The DMD resolution in our system is 1024×768 and the envelope size is set at 48×36 mm. A linear stage from VELMEX Inc (Bloomfield, NY) is used as the elevator for driving the platform in the Z axis. A rotary table also from VELMEX Inc is used to rotate the resin vats and cleaning stations. A high performance 4-axis motion control board with 28 Bi-directional I/O pins from Dynomotion Inc. (Calabasas, CA) was used for driving the linear stages and controlling the ultrasound cleaner, the fan, and a shutter. Two supper flat and clear glass Petri dishes are used as resin tanks. A PDMS film (Sylgard 184, Dow Corning) is coated on each glass dish. Figure 11.c shows the building of a test part in two different materials.6.2 Software System Figure 12 shows the flowchart of the multi-material MIP-SL process. Note the part will only be cleaned during the transitions between building in different materials. We also alternate the building sequence of two materials in neighboring layers (i.e.

Note the part will only be cleaned during the transitions between building in different materials. We also alternate the building sequence of two materials in neighboring layers (i.e. Ai → Bi → Bi+1 → Ai+1→…) such that less material switchover will be needed. A related multi-material MIP-SL software system has been developed by using the C++ programming language with Microsoft Visual C++ compiler. The graphical user interface (GUI) of the developed software system is shown in Figure 12.

The graphical user interface (GUI) of the developed software system is shown in Figure 12. The software system can synchronize the image projection and motion control based on geometry processing.prepare the motion and imagesGeometry slicingGeometry 1Geometry 2build base using single material processControl panelENDYfinished building?NMotion parametersonly Acurrent layeronly BA + BBuilding process reportbuild Abuild Abuild Bonly Bnext layerwash A and dry AA or A+Bnext layeronly BA or A+Bbuild Bclean A and dry AA or A+Bnext layeronly Bclean B and dry Bdock at CBdock at CAclean B and Dry Bdock at CBdock at CBdock at CAdock at CAFigure 12. Flow chart of the multi-materials MIP-SL system and a related software system.6.3 MaterialsTo test the multi-material MIP-SL system, we used Perfatory™ SI500 (yellow) and Perfactory™ Acryl R5 (red) from EnvisionTec Inc.

(Ferndale, MI), and FTI-GN (white) from 3D Systems Inc. (Rock Hill, SC). Beside the commercial resins, we also used a transparent resin that was developed inhouse. All thetested resins belong to acrylic resin. For the curing depth set at 0.1mm, the exposure time based on our projection system is set as 400 ms, 350 ms, 250 ms, and 300 ms for Perfatory™ SI500, Perfactory™ Acryl R5, FTI-GN, and transparent resins, respectively. Investigation on the bonding strength between the four different materials demonstrates that they can be bonded quite well due to the strong affinity between them.7. Results and Discussion A set of test cases have been designed to verify the developed prototype system in fabricating objects with different combinations of multiple materials. The experimental results have demonstrated that the presented two-channel bottom-up projection based approach can successfully build parts with desired material distributions.

The experimental results have demonstrated that the presented two-channel bottom-up projection based approach can successfully build parts with desired material distributions. Although the developed prototype system can only use up to two different materials, it can be extended in a straight forward manner to fabricate objects with three or more materials. 7.1 Components with Two Different Materials A main purpose of using multiple materials in a component is to provide additional functionality in the built part, such as varying colors, electrical conductivity, or mechanical properties.Verification on building objects with different colors: A test case based on the famous symbol of bagua is used to verify the bonding between two different materials (yellow and red resins) and to demonstrate the capability of the prototype system in building objects with different colors. The designed CAD model is shown in Figure 13.a.

The designed CAD model is shown in Figure 13.a. Accordingly the built object is shown in Figure 13.b.(a) CAD model(b) Built partFigure 13. A test case for verifying an object with different colors.40mmMask AMask BBuilt partFigure 14. A test case for verifying an object with different electrical conductivities.Verification on building objects with different electrical conductivities: Although most of acrylic or epoxy resins are electrically insulating, with proper modification, the resins could have good electrical conductivity (for example silver-filled epoxies). Embedding electrical circuits inside a 3D part is very meaningful for electrical and electronic design. In this way, the circuit with different shapes and differentorientation can be achieved (e.g. 3D circuit). In addition, the circuit can be designed to be adaptive to the target object shape (e.g. curved surfaces). A designed test case is shown in Figure 14.

3D circuit). In addition, the circuit can be designed to be adaptive to the target object shape (e.g. curved surfaces). A designed test case is shown in Figure 14. Masks A and B are the projection image for the conductive and insulative materials respectively (red and transparent resins). The built part is also shown in the figure, which verifies that the proposed method can be successfully used in the application.Verification on building objects with different mechanical properties: Another typical application of using multiple materials is improving the mechanical properties of designed components. For example, some portions of a product component may be soft while others are rigid. For the purpose, a designed test case is shown in Figure 15.a. The brush is mainly composed of two portions: the base and the brush head. They have different flexibility requirements, i.e. the base needs to be rigid while the brush head needs to be soft and flexible.

They have different flexibility requirements, i.e. the base needs to be rigid while the brush head needs to be soft and flexible. Two different materials (white and red resins) that have different mechanical properties are used in building the designed brush. The built object is shown in Figure 15.b. In addition, a USC’s Trojans logo with red material is embedded inside the white base. Based on the built part, we verified that the mechanical performances of the base and the brush head are different.5mm(a) CAD model(b) Built partFigure 15. A test case for verifying an object with different mechanical properties.4×48×816×1632×32Mask AMask BBuilt partFigure 16. A test case for verifying an object with digital materials.7.2 Components with Digital MaterialsAs demonstrated by the OBJET Connex family, the most unique feature of digital materials is that two base materials can be combined in specific concentrations and structures.

Therefore product components can have desired properties that may be different from those of the base materials. Accordingly, a designed test case is shown in Figure 16. A four grid slab with four different combination of two materials (red and yellow resins) is shown. The ratios of the two materials are 100% vs. 0%, 75% vs. 25%, 25% vs. 75%, and 0% vs. 100%, respectively. Using a halftoning method called dithering [22], we can get different combination of two materials by applying different dithering matrix. The built objects are also shown in the figure. The results demonstrate that our prototype system can mix two materials in predefined proportions for producing isotropic materials with different material properties.7.3. Limitations and ChallengesWe have presented a multi-material MIP-SL system for fabricating 3D components with spatially controlled digital materials. Multiple test cases have been performed to verify the presented approach.

Multiple test cases have been performed to verify the presented approach. In our tests, several issues and potential challenges have also been identified that need to be addressed in future work.Building speed: It takes ~3 minutes for the prototype system to finish the rough and final cleaning and drying. Such cleaning procedures take the majority of the cycle time. Better hardware design and new cleaning strategy may be explored to improve the building speed. • Material Waste: Even with the shallow vat design in our prototype system, we estimate there is still ~30% material waste due to the cleaning procedure.Material Bonding: The materials we tested all belong to the same type of resins. We have not tested the bonding between different types of resins (e.g. acrylic resins and conductive polymers).

We have not tested the bonding between different types of resins (e.g. acrylic resins and conductive polymers). The bonding strength between different types of materials may be an issue for future development on the multi-material MIP-SL process.Trapped air and bubbles: It is noticed that, for some types of geometries (e.g. holes and cavities), there are small bubbles in the built part. This is due to the trapped air in such geometries, which needs to be addressed in future research.Accuracy and resolution: To build digital materials, it is critical to accurately control the cured shapes with high resolution. Otherwise, the desired material combination ratio will not be achieved. Techniques such as the optimized mask image planning [23] can be used in improving the accuracy and resolution of cured shapes.8. ConclusionsA novel mask-image-projection-based stereolithography process has been presented for fabricating objects with digital materials.

ConclusionsA novel mask-image-projection-based stereolithography process has been presented for fabricating objects with digital materials. The proposed approach is based on projecting mask images bottom-up. Hence very shallow vat can be used in the building process. A new two-channel system design has been presented, which can significantly reduce the separation force of a cured layer from the resin vat. A two- stage cleaning strategy has been developed to avoid contamination during changing the platform between different resin vats. The fabrication results demonstrate that the developed dual-material MIP-SL system can successfully produce 3D objects with spatial control over placement of both material and structure.

The fabrication results demonstrate that the developed dual-material MIP-SL system can successfully produce 3D objects with spatial control over placement of both material and structure. The approach is general and can be easily extended from dual materials to multiple materials.The concept of digital materials as demonstrated by the polyjet process is interesting and significant.Our work illustrates that such a concept can also be achieved by other additive manufacturing processes for more selections on base materials. Combining multiple materials with various concentrations and structures to achieve desired characteristics such as multiple mechanical, electrical, chemical, biological, optical properties can have enumerable applications in the future. Hence the development of 3D multi- material printers is critical for future additive manufacturing research.References1.

viewed PaperDEVELOPMENT OF A NOVEL TEST ARTEFACT FOR CONFORMAL MATERIAL EXTRUSION PRINTINGSagar S Jalui1, Seyed Hossein Zargar2, Sheila Moroney1, Marcus Putz1, Mychal Taylor1, Serah Hatch1, Guha Manogharan1*1 Department of Mechanical Engineering, Pennsylvania State University, University Park, PA 16801 2 Department of Architectural Engineering, Pennsylvania Stata University, University Park, PA 16801Corresponding author: gum53@psu.eduAbstractAdditive manufacturing (AM) allows for free complexity. However, the layer-by-layer manufacturing method traditionally relies on a G-code input to the machine, representing 2D planar slices of each layer, which eventually combines to represent the net-shape 3D geometry. Through modification of existing slicer software, thus modifying the G-code input to the machine, non-planar (conformal) shells can be generated on top of a traditional planar scaffolding.

Through modification of existing slicer software, thus modifying the G-code input to the machine, non-planar (conformal) shells can be generated on top of a traditional planar scaffolding. The objective of this work is to design a novel test artifact to aid in the creation of design rules and to identify machine limitations for conformal printing. With the use of non-conventional design features using trigonometric (sine) surfaces, this test artifact would allow for deeper insights into the print quality of organic shapes made possible using a commercial, low-cost, material extrusion 3D printer. It would also enable the creation of design rules for conformal printing to push forward the true dual-Design for Additive Manufacturing (dual-DfAM) potential.Keywords: Conformal printing, test artifact, dual-DfAM, material extrusion, parametric design1.

IntroductionAdditive Manufacturing (AM) processes offer many design advantages and opportunities compared to traditional manufacturing processes. AM offers free complexity and one of its main components is shape complexity, indicating that AM can create geometrically complex shapes. However, in Material Extrusion (ME), there are issues with surface finish along organic curves as a result of the stair-stepping effect. However, non-planar ME printing can account for these issues and increase the shape complexity opportunities within desktop ME machines.Unlike traditional manufacturing techniques, Additive Manufacturing works by adding material layer by layer. By varying how each layer is printed very complex geometries can be created. For ME printing this is done by having a print head melt and then extrude out the plastic in the desired pattern [1]. However, there are some limitations that occur during slicing a geometry layerwise.

However, there are some limitations that occur during slicing a geometry layerwise. For example, as seen in Figure 1, when software is used to create the slices there is some information that is lost. In the figure this shows up as the staircase effect.1859Figure 1: Traditional method of generating planar slices from CAD [1]There are solutions to mitigate this problem which include decreasing the layer height, post-processing the part using acetone vapor smoothing, or using conformal printing. However, decreasing the layer height adds to the total build time and does not eliminate staircasing effect for high curvature geometries. Using post-processing steps leads to added costs as well as the need to compensate for the material that is smoothed during the operation.Several researchers have presented the use of conformal printing for various applications. Adams et al. [2] fabricated electrically small antenna on three dimensional surfaces using conformal printing.

Adams et al. [2] fabricated electrically small antenna on three dimensional surfaces using conformal printing. Aerosol jet printing has been utilized to fabricate conformal layers for applications in electronics packaging [3][4], thermoelectric devices [5], antenna [6], and sensors [7]. Shembekar et al. [8], [9] have also developed and optimized trajectories for conformal printing for robotic arms to enable large scale printing. However, most of these studies were performed using complex equipment or large-scale applications.Most of the complete and implemental research done on non-planar printing using desktop ME 3D printers was done by Daniel Ahlers’ master thesis [10]. The research focuses on generating G-code for conformal printing layers while accounting for various machine-nozzle configurations.

The research focuses on generating G-code for conformal printing layers while accounting for various machine-nozzle configurations. The output of this research provides a modified version of Slic3r, a free slicing software, that allows user-friendly access to the necessary parameters required to create a successful print. As shown in Figure 2, the modified version of Slic3r also shows the user the parts of the print which will be printed using conformal layers, shown in the pink with the planar layers shown in yellow.Figure 2: Example of an output from the Slic3r modification for non-planar slicing [10]1860The nozzle geometry is a critical parameter and must be measured for the input to the slicer to allow it to generate the correct G-code.

The important geometry of the nozzle is the nozzle height and nozzle angle as shown in Figure 3.Figure 3: Illustration of two critical parameters from a commercial desktop material extrusion 3D printer while generating non- planar layers [10]These dimensions are important as the angle of the nozzle and the height of the nozzle determines how steep of an angle the non-planar part can be before the nozzle collides with the previous layer.Another potential solution to enable conformal printing using desktop 3D printer modification was presented by Hong et al. [11]. They proposed a method for the modification of a Prusa i3 to allow 5-axis 3D printing. An example of this setup can be seen in Figure 4.Figure 4: Example of Prusa i3 modification for 5-axis printing [11]As can be seen from the different examples of conformal printing the designer has more variables that they need to consider when first creating a part and when preparing the part for printing.

As listed above, the main variables that need to be considered are the dimensions of the nozzle on the printer and the overall geometry of what is printed. The geometry of the part needs1861to be designed such that it takes advantage of conformal printing, smooth surface finish, stronger parts, and provides an avenue for opportunistic DfAM features.In previous work, researchers at the National Institute of Standards and Technology (NIST) created a test artifact to evaluate how well a specific metal AM machine can handle various kinds of design features [12]. As seen in Figure 5 below, the test artifact includes a large variety of design features. There are holes and pins that test both repetition and minimum feature size. There are also “staircases” and a ramp to test other features that are commonly seen in metal AM parts. Each feature can be quantitatively measured to easily identify how well the machine has produced them.

Each feature can be quantitatively measured to easily identify how well the machine has produced them. Each of these features falls under the category of restrictive Design for Additive Manufacturing (DfAM). If a machine cannot complete a certain feature, then limits can be set for future design of parts. Knowing these limits before building an important part is very valuable. There is a lot of room for future work in this area, in creating test artifacts for other AM processes outside of metal AM.Figure 5: NIST specimen; The most popular standard AM test artifact [12]The objective of this paper is to propose a novel test artifact design for conformal printing. This artifact contains contemporary parametric design features to mathematically represent and control the freeform geometry of non-planar surfaces. Section 2 expands on preliminary research conducted using a sample geometry.

Section 2 expands on preliminary research conducted using a sample geometry. Section 3 highlights the fundamental parametric design features and the proposed test artifact design. Section 4 discusses the potential design of experiments and analysis for the proposed test artifact design.2. Preliminary ExperimentationTo understand the working of the Slic3r modification, a virtual machine was installed to run Ubuntu, the preferred environment to compile the graphical user interface (GUI) of the Slice3r modification. The GUI contained a separate section for activating non-planar slicing.

The GUI contained a separate section for activating non-planar slicing. The parameters along with their interpretations have been summarized in Table 1.1862Table 1: Table showing the list of user controllable parameters using Slic3r modified conformal slicing tool [10]Parameter Maximum nonplanar angleMaximum nonplanar collision angleMinimum nonplanar areaMaximum nonplanar collision heightIgnore collision sizeInterpretation Sets the maximum angle where the surfaces are printed nonplanar The maximum angle that can be printed without collision. This is only used for collision checking. The setting depends on the printer. 90° Disables collision checking Throws away all regions that are smaller than this to avoid very small nonplanar areas. Depends on the printed model. Sets the maximum height of nonplanar surfaces. Maximum nonplanar collision angle must be collision free in this region. The also depends on the hardware you use. Set to 999999 to disable.

Sets the maximum height of nonplanar surfaces. Maximum nonplanar collision angle must be collision free in this region. The also depends on the hardware you use. Set to 999999 to disable. Ignores collisions smaller than this value. Basically, a workaround to get rid of small artifacts generated by the offsetting algorithm that is used for collision checking.Using the parameters in a stock Prusa i3MKS printer, a half dome with a diameter of 1.5 inches and a height of 0.25 inches was sliced using the non-planar slicing settings. The maximum height was set at 7.5 mm and the maximum angle was set at 45°. One of the major challenges we faced was that the probe for the auto-bed leveling sensor was obstructing the part geometry during the non-planar layers’ printing. Thus, a stock Prusa mini was chosen for testing purposes. The half dome was printed using standard PLA 1.75 mm diameter filament with a 0.4 mm nozzle maintained at 215°C, while the bed temperature was maintained at 60°C.

The half dome was printed using standard PLA 1.75 mm diameter filament with a 0.4 mm nozzle maintained at 215°C, while the bed temperature was maintained at 60°C. The total print time was 5 minutes 13 seconds. The final printed part can be seen in Figure 6.Figure 6: The test half-domes printed using conformal layers on the Prusa mini1863The first learning was that the non-planar layers were printed on top of a planar base. The slicer generated planar layers to approximate the geometry and then printed the last few layers of the shells using non-planar material deposition. Additionally, it was noted that there was filament accumulation at the corner where the non-planar deposition started and ended between layers. Overall, it was observed that the final quality was visually more appealing as compared with the standard planar slicing.

Overall, it was observed that the final quality was visually more appealing as compared with the standard planar slicing. The conformal layers resulted in a much smoother surface.It is also important to note that these tests were performed using a stock Prusa mini printer. There are several nozzle modifications available to increase the clearance between the nozzle and the extruder to increase the “maximum non-planar height” and “maximum non-planar angle” parameters that allow for varied regions of conformal slicing. One such modification by “nonplanar.xyz” [13] can be seen in Figure 7.Figure 7: Example of nozzle modification to enhance non-planar printability [13]3. Parametric Design of SpecimenDue to the various range of variables available for assessing the functionality of a designed non-planar specimen, different approaches were considered for parametric representation of variables.

The initial phase of the design began by defining different parameters related to non- planar slicing of the final specimen that have the highest potential of assessing the accuracy of the printing process. These parameters are considered based on existing literature and were divided into three different sub-categories. As shown in Figure 8, the parametric ribbon (sinusoidal) will focus on the highest and lowest points to reach in a non-planar scenario without failure. The variables are defined in a 2D interface based on a gradually changing sine function, and the dimensions of the final specimen can be modified as well. It is worth mentioning that the parametric workflow will give the designers and engineers the opportunity to adjust the design variables interactively based on specific characteristics of AM manufacturability. In general, based on a printer’s assembly and requirements, the final geometry can be altered.

In general, based on a printer’s assembly and requirements, the final geometry can be altered. The second part of Figure 8, focused on gradually changing the curvature in 3D space. By focusing on freeform hyperbolic form, the end effector’s range of motion for printing a double curvature geometry is investigated. The hyperbolic form is designed parametrically using Python coding and changing the variables will generate a variety of forms. The final range of variables is included in designing1864a two-curvature dome-shape geometry. The variables will generate a sub-specimen with gradual changing of curvature in a 3D space sing force density form-finding method. The final geometry will assess the possibility of double-curvature printing in a more controlled environment.

The final geometry will assess the possibility of double-curvature printing in a more controlled environment. By changing the parameters related to the form finding process, the highest peak point can be determined to assess non-planner criteria.Figure 8: Different parametric prototypes using Rhino/GrasshopperFocusing on the three different parametric geometries, a final ‘parametric workflow’ was defined to generate the needed specimen. The workflow was based on the initial sinusoidal function that gives the designer the opportunity to change amplitude and frequency to generate design alternatives. By gradual change of height to the base of the specimen, mountain and valley- like series of shapes was extracted. The geometry was blended to provide a baseplate for printing and the final geometry was exported as an STL file. Figure 9 illustrates the range of variables including thickness of each part.

The geometry was blended to provide a baseplate for printing and the final geometry was exported as an STL file. Figure 9 illustrates the range of variables including thickness of each part. For the proposed test artifact design, nine different case studies are generated by a combination of amplitude and frequency in low, medium, and high settings for a full factorial design. The final geometry can be replicated as many times as needed to perform accurate data-based printed design analysis. As shown in Figure 9, different models can be merged to generate a single specimen ready to perform the slicing process. Having low, medium, and high settings will provide a reach dataset to assess the importance of frequency and amplitude of initial sine function on slicing and actual printing process.1865Figure 9: Final prototype of the proposed test artifact with variables related to parametric workflow4.

Design of Experiments for Conformal Test ArtifactUsing the learnings from the preliminary experimentation, and the proposed specimen design, the proposed design of experiments contains the full factorial analysis of two design variables varying over three levels. To identify the limit of generating conformal layers, the Slic3r modification GUI was used. However, to include a more comprehensive analysis it is proposed that the conformal sliced printed geometry must be compared to the planar sliced printed geometry. Both parts will be printed using the same layer thickness using three replicates per slicing method to account for variation. Post-printing, the parts will be 3D scanned using the LiDAR scanners present in a modern-day smartphone. The failure criteria would be guided by the root mean square (RMS) surface deviation between the measured data (3D point cloud obtained using the 3D scanner app), and the CAD surface data.

The failure criteria would be guided by the root mean square (RMS) surface deviation between the measured data (3D point cloud obtained using the 3D scanner app), and the CAD surface data. Such analysis can be done using the advanced inspection software Geomagic Control X by 3D Systems. The RMS difference would account for a surface averaged quantifiable measure of surface quality. If the RMS difference between the measured and design data is within 10% of the layer thickness, it will be considered a pass, otherwise, it will be declared a failure point. The failure point will be used to identify areas where conformal printing would be beneficial.The results obtained from Slic3r using non-planar settings for the parametric features as well as the final test artifact have been shown in Figure 10 and Figure 11 respectively. The non- planar regions can be seen in pink whereas the planar regions are represented in yellow.

The non- planar regions can be seen in pink whereas the planar regions are represented in yellow. It was also observed that by changing the nozzle configurations, different regions of conformal printing were possible. It was interesting to note that in the test artifact, there were substantially low non- planar layers generated within the negative sinusoidal features. This was attributed to the low “maximum nozzle height” parameter which makes it difficult to print depressions while being able1866to print peaks. It was observed that low amplitudes tend to be easier to fabricate using the stock hardware in desktop 3D printers. It was also noted that lower frequencies tend to be favored while using non-planar printing.Figure 10: The different parametric features sliced using planar and non-planar slicingFigure 11: The proposed test specimen design sliced using non-planar slicing (left); Close up of the non-planar layers (right)18675.

ConclusionThis paper proposed a novel test artifact design for conformal printing. The designing process contained parametric design features such as sinusoidal, hyperbolic, and double curvature surfaces that represent a contemporary approach to rethinking design features for organic looking geometry possible using AM. The final test artifact contained two design variables – amplitude, and frequency, at 3-levels for a full factorial design. The test artifact was sliced using the non- planar settings within the Slic3r modified GUI using printer parameters for a stock material extrusion desktop 3D printer. It was proposed that three replicates of the test artifact be printed using planar and non-planar techniques. These parts would be compared on the basis of RMS surface deviation between measured geometry and as-designed geometry using Geomagic Control X.

These parts would be compared on the basis of RMS surface deviation between measured geometry and as-designed geometry using Geomagic Control X. Additionally, the limits of generating non-planar layers using the traditional desktop 3D printer would be identifiable from the slicer using multiple machine-nozzle configurations.The overall objective of this study is to answer the question, “If it can be printed conformally, does it mean we should print it?” The conformal layers require a planar set of layers to create a base upon which it adds conformal layers. Right now, these planar layers represent solid material, but it could be modified to represent removable support material which would provide avenues for generating freeform “shell” geometries that could be used for potential applications where surface quality is of utmost importance (e.g. injection molding, conformal cooling channels).

injection molding, conformal cooling channels). However, surface quality can further be enhanced using post-processing steps such as acetone vapor treatment for PLA parts. It is possible that future studies might show enhancement in cycle times and quality for such post processing methods using conformal layers. Authors are currently continuing experimental investigations into qualitative and quantitative comparisons between the test artifact printed both using planar and conformal slicing. Limitations of this study include the use of only the existing Slic3r modification software to generate planar layers, comparing design limits at three levels using two non-conventional design features. The work remains conceptual at this stage with future work being the fabrication of the test artifact using stock, and modified nozzles.

The work remains conceptual at this stage with future work being the fabrication of the test artifact using stock, and modified nozzles. The future work also includes the addition of a wider range of parametric design features to expand the possibilities using conformal printing.AcknowledgementsThe authors would like to thank the instructor Dr. Nicholas Meisel and students of the class for their valuable input through the peer-review assignments in the graduate-level course on design for additive manufacturing.References[1]G. A. Nisja, A. Cao, and C. Gao, “Short review of nonplanar fused deposition modeling printing,” Mater. Des. Process. Commun., vol. 3, no. 4, p. e221, Aug. 2021.[2]J. J. Adams et al., “Conformal Printing of Electrically Small Antennas on Three- Dimensional Surfaces,” Adv. Mater., vol. 23, no. 11, pp. 1335–1340, Mar. 2011.[3]J. A. Paulsen, M. Renn, K. Christenson, and R.

VELOPMENT OF A PRECISION RAPID METAL FORMING PROCESSF.W. Liou*, J. Zhang*, S. Agarwal+, J. Laeng*, and J. Stewart*Department of Mechanical and Aerospace Engineering and Engineering Mechanics, University of Missouri-Rolla, Rolla, MO 65409 + Department of Electrical and Computer Engineering, University of Missouri-Rolla, Rolla, MO 65409 AbstractThis paper presents the important issues pertaining to the development of a precision rapid metal forming process. A five-axis configuration provides a flexible building capability to produce free-form fabrication capability. The laser cladding process is able to produce functional mechanical parts and machining capability is able to produce industrial grade surface quality. A machine configuration that combines the laser cladding and CNC machining processes is presented.

A machine configuration that combines the laser cladding and CNC machining processes is presented. The related parameters and components are discussed.IntroductionDue to the increasing complexity of certain part geometry, producing highly accurate functional models directly from computerized data is often very difficult using the available rapid prototyping (RP) systems. The laser metal forming process is a potential technique that can solve this problem. LENS (Laser-Engineered Net Shaping) is one developing Rapid Metal Forming (RMF) process that has demonstrated the feasibility of laser metal forming to produce near-net-shape metal parts [2, 3]. Some efforts have been dedicated to utilizing the existing laser cladding technique, which is an established method in surface coating of metal parts, to produce metal parts by a layer additive approach [1, 4]. The laser cladding process is able to produce fully functional mechanical parts.

The laser cladding process is able to produce fully functional mechanical parts. However, these parts usually pose inadequate accuracy compared to that of the CNC machining process. To address such issues, a machine configuration that combines the laser cladding and CNC machining processes is presented in this paper.Rapid Metal Forming Process (RMF)Rapid metal forming systems, especially those of laser-based energy sources, were reported by many authors [1-6]. These systems represent some current efforts to produce metal parts directly from CAD solid models. A number of studies on rapid metal forming with lasers were primarily concerned with the process parameters that determine the quality of the produced parts [1, 4, 6].

A number of studies on rapid metal forming with lasers were primarily concerned with the process parameters that determine the quality of the produced parts [1, 4, 6]. Most of these systems are still in the research and development stage.362The basic system hardware for a rapid metal forming process includes an energy delivery system (laser head, optics system), powder delivery apparatus (powder feeder, nozzle, carriage gas), and a Computer Numerical Control (CNC) table for x-y motion and z-axis vertical motion. The CNC table can have three-axis or multi-axis motion for enhanced flexibility that is needed in forming more complex parts [1, 2, 4-8]. Figure 1 shows the basic system hardware required for a metal forming system.Reflective Mirror directing Laser beam onto focusing lenseLaser BeamLaser SourceFocusing LenseLaser and Powder nozzleBase PlateLaser SpotCNC TableFigure 1.

Basic System Hardware Required for a Metal Forming SystemIn addition to the basic hardware, a metal forming system also consists of control system hardware, a cooling system for the powder nozzle, and a base plate. Control system hardware includes sensors, CCD cameras, and a computer workstation. Online control is an integral part of a metal forming system. A cooling system is used for the powder nozzle because it is exposed to a thermal load by scattered and reflected laser radiation [4, 6, 9]. A base plate is normally used as a substrate on which metal layers are deposited until a complete part is formed. The base plate is removed by machining or by dissolving with certain chemicals, leaving the finished metal part.363Making Complex Part GeometryA robust metal forming system should be able to create prototypes of any complex geometry that requires deposition at inclined angles or on curved surfaces.

For example, a three- axis system cannot build multi-layer walls with 30(cid:176) angles from the vertical plane as illustrated in Figure 2, and therefore, a more flexible system, such as a five or six axis machine, is required.Support structure should also be minimized. Building in the vertical direction, the system cannot build a “T” shape structure without excessive support structure. However, a greater degree of freedom system with rotational axes in both the x and y directions, can build overhang geometry with minimum support structures needed.Build DirectionMetal LayerVertical Plane30Build DirectionVertical WallFigure 2. Multi-layer Walls at 30(cid:176) Incline364Dimensional Accuracy ConsiderationIn a report on laser metal deposition system [10], a sample, as shown in Figure 3, is used to obtain a perspective of the dimensional accuracy of parts in the x, y, and z directions.

Measurements taken in the x and y directions are within 0.002 inches (0.05mm) of the nominal dimension. The accuracy of the z dimension is typically within 0.015 inches (0.38mm) from nominal. The accuracy of the z direction is dependent on the build layer thickness. The thinner the build layer, the more accurate the part. The ability to build a part with variable layer thickness would improve the accuracy in the z direction. The surface roughness of the part is relatively rough compared to a machined surface. Surface roughness is a function of powder particle size, and the optimum particle size for the process produces a surface between 200-300 microinch average roughness (Ra). Therefore, in comparison with the current laser cladding processes, the machining process is able to produce parts with superior accuracy and surface finish with a broader range of materials.Figure 3.

Geometry of the Sample in the Accuracy Testing [10]LAMP Laboratory ProcessThere are some general challenges in the current SFF (Surface Free-from Fabrication) processes, such as to produce parts with the accuracy and surface finish required, to build parts with a wide variety of engineering materials, and to directly produce high-quality metal parts for production tooling applications. The automated CNC machining process can also be used as a rapid prototyping tool. Current CNC systems, however, are not extensively used for SFF365because of the reasons such as the facts that skillful human intervention required to plan the operations and to operate the equipment, custom fixturing and special tooling often required, and inherent geometric limitations, as shown in Figure 4, hindering the process flexibility.Figure 4 Geometric Limitation of the CNC Machining ProcessesA combined system may have the advantages of both laser cladding and machining.

It allows complex parts to be built with high accuracy. The schematic of the combined laser-based metal forming LAMP system is shown in Figure 5. The system includes two major subsystems: a laser deposition system and a CNC milling machine system. These two systems share the common xy translator and rotary-tilting table; however, they have independent z movement. The laser deposition system and CNC milling machine work in shift in a five-axis motion mode. The laser deposition system is divided into two independent systems: a laser system and a powder feeder. The laser system includes a 2,500W Nd-YAG laser, a laser beam delivery module, a focus optics module, a five-axis movement module, an optics protection module, etc. The laser system has a combined movement configuration, with z-motion from the motion of the optics, xy translation, and rotating and tilting of the workpiece.

The laser system has a combined movement configuration, with z-motion from the motion of the optics, xy translation, and rotating and tilting of the workpiece. The other important subsystem of laser deposition is the powder feeder, which feeds powder into the melting pool formed by the laser beam.366Figure 5. Schematic of the 5-Axis Material Deposition/Removal SystemConclusionThe LAMP system which features both material deposition and material removal processes and 5-axis capability will have the advantages of quickly producing functional parts with precision surface.

The expected advantages of this process include 1) the capability to build complex parts including free form surfaces and thin walls; 2) much improved surface finish which is essential in tooling applications; 3) reduced structure support in building 3-D objects that is critical in making metal/ceramic parts; 4) Opportunities of reduced build time since thicker layers can be deposited and shaped by a machining process; 5) almost no limitation on material selection as long as the laser can melt the material in time; 6) capability of making perfect features such as holes and threads using machining processes; 7) Elimination of shrinkage factor in deposition process with the material removal process. This is not only useful in modern manufacturing environment in producing functional prototypes but also producing molds for mass production.

This is not only useful in modern manufacturing environment in producing functional prototypes but also producing molds for mass production. The challenges, however, include the complexity in process planning and the system control.AcknowledgementThis research was supported by the National Science Foundation Grant Number DMI-9871185, Missouri Research Board, and a grant from the Missouri Department of Economic Development through the MRTC grant. Their support is appreciated.367References1.2.3.4.5.Koch, J.L. and Mazumder, J., “Rapid Prototyping by Laser Cladding,” The International Society for Optical Engineering, vol. 2306, 1993, p. 556-eoa. Keicher, D.M., Miller,W.D., Smugeresky, J.E. and Romero, J.A., “Laser Engineered Net Shaping (LENS): Beyond Rapid Prototyping to Direct Fabrication” , Proceeding of the 1998 TMS Annual Meeting, San Antonio, Texas, USA, p.

VELOPMENT OF A PROTOTYPE LOW-VOLTAGE ELECTRON BEAM FREEFORM FABRICATION SYSTEMJ.K. Watson*, K.M. Taminger†, R.A. Hafley†, and D.D. Petersen*NASA Johnson Space Center, Houston, TX 77058 † NASA Langley Research Center, Hampton, VA 23681AbstractNASA’s Langley Research Center and Johnson Space Center are developing a solid freeform fabrication system utilizing an electron beam energy source and wire feedstock. This system will serve as a testbed for exploring the influence of gravitational acceleration on the deposition process and will be a simplified prototype for future systems that may be deployed during long-duration space missions for assembly, fabrication, and production of structural and mechanical replacement components. Critical attributes for this system are compactness, minimal mass, efficiency in use of feedstock material, energy use efficiency, and safety.

Critical attributes for this system are compactness, minimal mass, efficiency in use of feedstock material, energy use efficiency, and safety. The use of a low-voltage (<15kV) electron beam energy source will reduce radiation so that massive shielding is not required to protect adjacent personnel. Feedstock efficiency will be optimized by use of wire, and energy use efficiency will be achieved by use of the electron beam energy source. This system will be evaluated in a microgravity environment using the NASA KC-135A aircraft.INTRODUCTIONFuture long-duration human exploration missions will be challenged by constraints on mass and volume allocations available for spare parts. Addressing this challenge will be critical to the success of these missions.

Addressing this challenge will be critical to the success of these missions. As a result, it is necessary to consider new approaches to spacecraft maintenance and repair that minimize the mass and stowage volume that must be allocated for spares while enhancing mission robustness [1].Production of replacement components by solid freeform fabrication (SFF) processes during a mission could reduce or eliminate the need to carry a complete inventory of pre- manufactured spares. Rather, replacement components would be generated as needed from feedstock material. As a result, only the total mass of replacements would need to be estimated instead of a prediction of which specific components might be needed. Attempting to predict which components will fail and require replacement will inherently be an inaccurate process and will certainly result in provisioning numerous components that will never be used (wasted mass) and may result in under-provisioning of other components.

This technology could also be used to support fabrication of large space structures and repair of spacecraft primary structure [2].Direct deposition processes utilizing laser or electron beam energy sources offer the greatest potential for space applications. Although the feedstock material can be introduced as either a powder or wire, the wire form is preferable for in-space application because of458operational and safety issues associated with management of metallic powder in a microgravity environment. The basic electron beam freeform fabrication (EB F3) process and preliminary results obtained in ground-based experiments are described in reference 3.Key issues associated with the production of components during space missions center on the potential effects of a microgravity environment on the process. In the microgravity environment, body forces on the molten region of the deposited bead become insignificant and surface tension forces will dominate.

In the microgravity environment, body forces on the molten region of the deposited bead become insignificant and surface tension forces will dominate. Therefore, the cross-sectional shape of the molten region may be different during microgravity deposition than during deposition in a 1-g environment. This will have at least two potentially significant effects. First, the bead cross-section will govern the topography of the exposed surfaces of the final part. The resulting surface condition and associated stress concentrations will affect the mechanical properties of the component. Second, closed-loop process control schemes may rely on monitoring the shape of the molten portion of the deposited bead. Since the bead shape in microgravity may be somewhat different from that in a 1-g environment, it is essential that the nominal shape in microgravity be fully characterized.

Since the bead shape in microgravity may be somewhat different from that in a 1-g environment, it is essential that the nominal shape in microgravity be fully characterized. Also, solidification of the material in the microgravity environment may affect metallurgical characteristics.For these reasons, an important initial step in developing an SFF process suitable for use in space will be an experimental effort to understand and define the effects of a reduced-gravity environment on the process. For the near term, work will be performed using the NASA Johnson Space Center (JSC) Reduced Gravity Program KC-135A research aircraft because of its accessibility and relatively low cost. As the process and apparatus mature it will be appropriate to take advantage of facilities such as the Space Shuttle or International Space Station. Compactness, low mass, and efficient energy utilization will be essential attributes for SFF system hardware that will be used in any of these facilities.

Compactness, low mass, and efficient energy utilization will be essential attributes for SFF system hardware that will be used in any of these facilities. Particularly because of superior energy conversion efficiency compared with laser systems, electron beam technology was chosen for the current effort. This characteristic is an attribute both for the current development system and for future operational systems that may evolve from this work.The initial phase of the effort will consist of system development and ground-based checkout, initial KC-135A microgravity flight tests, and sample analysis.

The primary goals of this phase will be to: (1) develop a working prototype EB F3 system for reduced gravity flight, and (2) evaluate of the effects of a microgravity environment on the metallurgy and geometry of the material deposited with this process.EXPERIMENTAL SYSTEM DESIGNAn artist’s conception and a schematic diagram of the low-voltage EB F3 system being designed for microgravity flight tests are shown in Figures 1 and 2, respectively. The major components that will be contained within the vacuum chamber include the electron beam gun, positioning system, and wire feeder. The electron beam gun and wire feeder will be mounted directly to the top portion of the vacuum chamber. The positioning system and ports for the vacuum system will be mounted to a base that can be dropped down to provide full access to the interior of the chamber when not under vacuum.

This hazard is controlled by incorporating sufficient lead shielding in the walls and window of the vacuum chamber. Clearly, massive shielding would be counter to the objective of developing a lightweight, portable system for use on research aircraft. Further, the mass associated with this shielding would be a very undesirable characteristic for an operational system to be used onboard spacecraft. Since heavy shielding is undesirable in this application, the alternative approach is to design the process in such a way that production of hazardous penetrating radiation is avoided. This can be accomplished by operating at significantly lower accelerating voltages than are typically used for industrial electron beam welding. Operation in the range of 8 – 15 kV will minimize production of penetrating radiation [4].

Operation in the range of 8 – 15 kV will minimize production of penetrating radiation [4]. Shielding provided by a simple stainless steel or aluminum vacuum chamber should be adequate to ensure the safety of operating personnel [5].The electron beam gun will generate a focusable low-voltage beam with a maximum beam power of 3-5 kW and will operate using 110 V input power. This design was selected to provide sufficient power density for the EB F3 process and still meet safety considerations for radiation shielding. The gun is compact in size and operates using the vacuum from the process environment, thereby eliminating the requirement for auxiliary vacuum pumping of the gun. The filament will be selected for robustness and long life.

The filament will be selected for robustness and long life. As shown in the schematic in Figure 2, the electron beam system is designed for an electrically conductive path to ground the system to the aircraft.Vacuum SystemThe electron beam deposition process must be performed in a vacuum to prevent attenuation of the electron beam and ionization of gases in the process environment. This also provides a clean atmosphere to prevent contamination of the metal during the deposition process. The vacuum level needs to be a minimum of 10-5 torr for proper operation of the electron beam gun. Size, weight, power, and rapid evacuation requirements restrict the choices of vacuum pumps used for the simulated microgravity experiments performed on the KC-135A. In addition, the vacuum system must be robust enough to withstand the cyclic low-g to high-g loading conditions plus any unpredictable loads experienced due to air turbulence.

In addition, the vacuum system must be robust enough to withstand the cyclic low-g to high-g loading conditions plus any unpredictable loads experienced due to air turbulence. (Further details on the flight conditions are described in the “Aircraft Operations and Safety Requirements” section below.) The final design for this system calls for three vacuum pumps: an oil-less scroll pump, a turbomolecular pump, and an ion pump. The scroll pump will achieve a rough vacuum, and serve as a backing pump for the turbomolecular pump. At approximately 1 torr, the turbomolecular pump will switch on to evacuate quickly down to the 10-6 torr range. Once the chamber reaches 10-6 torr, the valve to the turbomolecular pump will be closed and the pump shut off to protect the high speed rotating vanes from damage that may occur due to sudden loading or changes in gravitational forces on the pump.

The ion pump will then be turned on to maintain high vacuum and to remove any outgassing or metal vapor resulting from the EB F3 process. A further precautionary measure will be to mount the pumps with shock461suppression to protect them from damage due to irregular loading conditions as a result of the parabolic flight trajectories and unforeseen aerodynamic turbulence.In space, the equipment can be designed and operated to use the vacuum of space as the process environment for the EB F3 process. The processing equipment might either be located such that it can be fully exposed to the space environment or located in a working chamber that is vented overboard. In either case, use of the high vacuum readily available in space will significantly simplify the EB F3 system design by eliminating the vacuum system.Positioning SystemThe positioning system has been designed to move the substrate while allowing the electron beam to remain stationary.

This approach does not maximize efficiency of the build envelope, but it significantly simplifies the system design. The position where the wire is fed into the focal point of the electron beam is a fixed point. This allows for cameras to easily be positioned to monitor the process and provide feedback control. The positioning system has four axes: X, Y, and Z linear axes, plus a-axis rotation in the X-Y plane. During the weightless portions of the flight, the deposition will occur while translating in the X-Y plane using X and Y linear travel and/or rotation to build a single layer. During the high-g portions of the flight, the deposition process will be suspended and the substrate will be indexed down in the Z-direction a distance equal to the height of the previous layer, then repositioned in X, Y, and rotated back to a starting point in preparation for the next pass.

This initial approach will maintain a constant angle between the wire feed orientation and the translation direction. The four-axis system allows sufficient flexibility for more detailed process development and the ability to produce more complex parts in the future.The positioning system has been designed specifically for the EB F3 process. The precision and accuracy of the positioning system are specified to be smaller than the molten pool. The translation speed ranges are designed to allow a large process envelope in which control of the build height and width and deposition rates can be optimized for specific structural features and metallurgical characteristics. The mechanisms are covered to protect them from contamination with condensed metal vapor from the EB F3 process. In addition, the positioning system is thermally and electrically isolated from the process zone to protect it from the heat and electrical current generated by the electron beam.

In addition, the positioning system is thermally and electrically isolated from the process zone to protect it from the heat and electrical current generated by the electron beam. The motors are vacuum-compatible and contained within the chamber.Wire Feed SystemThe feedstock form selected for this process is metal wire. Powdered metal feedstock requires gravity or flowing gas to direct the powder into the molten pool. For operating in a vacuum at microgravity, neither of these powder delivery methods will work. In addition, containment and handling of powdered metal poses significant safety issues in a microgravity environment. The use of wire feedstock eliminates the need for flowing gas and provides nearly 100% feedstock usage efficiency, resulting in the minimum mass and virtually no waste products.

The use of wire feedstock eliminates the need for flowing gas and provides nearly 100% feedstock usage efficiency, resulting in the minimum mass and virtually no waste products. The wire feeder is capable of feeding very small diameter wires at both high and low speeds, enabling a range of deposition rates and fine detail to be achieved within this EB F3 system. The one-pound462wire spools and the wire feed system will be vacuum compatible for location within the vacuum system close to the process zone.InstrumentationThe instrumentation for the low-voltage EB F3 system is designed to measure processing conditions and record data throughout the processing, indexed on a common time scale. All of the controls and data acquisition will be programmed through a LabView interface on a PC laptop computer. The vacuum chamber will be equipped with multiple viewing ports for mounting high speed and standard video cameras to visually record the active process zone during microgravity.

The vacuum chamber will be equipped with multiple viewing ports for mounting high speed and standard video cameras to visually record the active process zone during microgravity. Accelerometers will be attached to the equipment to measure the gravitational forces and accelerations. In addition, beam parameters, vacuum levels, positioning, and wire feeder control parameters will be recorded on the same time base to match with the process monitoring instrumentation outputs. Additional electrical feedthroughs will be incorporated into the vacuum chamber wall to allow implementation of other monitoring devices, such as thermocouples, for future experiments.Aircraft Operations and Safety RequirementsTesting of the EB F3 system in a reduced gravity environment will be conducted aboard NASA’s KC-135A aircraft. Specially modified for the reduced gravity operations and maneuvers, the aircraft flies along a trajectory in a series of parabolas.

Specially modified for the reduced gravity operations and maneuvers, the aircraft flies along a trajectory in a series of parabolas. Weightlessness, from partial gravity to zero-g depending upon the specific trajectory flown, is experienced during the top portion of the parabola. Up to 1.8-g may be felt during the pull-out and bottom of the parabola, and then again during the initial pull-up (ascent). The expected time for reduced gravity is 20 to 23 seconds for each parabola, with a standard of 40 parabolas per flight.Because the time of reduced gravity is short-lived, the objectives of the experiment must be concise and well planned to be completed within the 20 second zero-g period; otherwise, the results of the research can be jeopardized by the affects of the 1.8-g pull-out.

In addition, many other requirements must be met, both for the test engineers (medical certification consisting of a flight-rated physical, and physiological training) and the flight experiments.Prior to flight a test equipment data package will be submitted and will include a test plan, engineering drawings, structural analysis, electrical loads analysis, and all identifiable hazards. The electron beam freeform fabrication system will operate on 110 volt AC, 60 Hz, single phase input power, one of four available power interfaces provided for aircraft payloads. The system will also be designed to handle the possible occurrence of brief power interruptions during flight.

The system will also be designed to handle the possible occurrence of brief power interruptions during flight. For contingency purposes, structural analysis will be performed to verify the experimental hardware and all major components will withstand the maximum, worst case, take- off and crash-landing loads of 9-g’s.As previously mentioned, a form of shock suppression will be used on the vacuum pumps to dampen the 1.8-g effects experienced during pull-out and ascent. However, all subsystems and components will require either test or analysis to verify their operation through the 1.8-g maneuvers, including components in the positioning system.

However, all subsystems and components will require either test or analysis to verify their operation through the 1.8-g maneuvers, including components in the positioning system. Other items to evaluate will463include any inherent risks of the vacuum system, and requirements for overboard venting of outgases and metal vapors resulting from the EB F3 process.PLANNED TEST PROGRAM AND SYSTEM EVOLUTIONThe low-voltage EB F3 system as described in this report will be tested on the ground prior to flight testing on board the KC-135A. This will serve to work out the test matrix and provide baseline parts constructed in a 1-g environment. Experiments conducted on the KC- 135A will be used to study the effects of microgravity on the geometry of the molten pool, the resulting build geometry, and solidification microstructure.

Experiments conducted on the KC- 135A will be used to study the effects of microgravity on the geometry of the molten pool, the resulting build geometry, and solidification microstructure. Different molten pool diameters will be explored to examine regions where the process is surface tension dominated (with a small diameter molten pool) as compared to being dominated by gravitational forces (with a larger diameter molten pool). Research with another ground-based EB F3 system at NASA Langley will focus on exploring process parameter domains and defining acceptable operating envelopes through correlation of processing, microstructure, and mechanical properties [3]. Processing parameters will be developed for a variety of alloys of interest for aerospace structures, including aluminum, aluminum-lithium, and titanium alloys.

Processing parameters will be developed for a variety of alloys of interest for aerospace structures, including aluminum, aluminum-lithium, and titanium alloys. In addition, comparing parts built using the low-voltage system with those built in the ground-based system at NASA LaRC will assess reproducibility and robustness of the EB F3 process.The low-voltage EB F3 system has been designed to allow for adding capabilities for enhancing the process controls. Such evolution may include addition of the ability to change the orientation of the wire feed angle into the molten pool, extended part manipulation capabilities, development of closed-loop controls and process automation, and the ability to build parts directly from computer-aided design (CAD) files.

The results of testing performed in 1-g and microgravity environments will assist in deciding the appropriate directions and capabilities to be developed in future efforts to enable the deployment and application of such a system in support of human exploration missions.SUMMARYElectron beam freeform fabrication (EB F3) is a new process being developed that has promise for supportability of long-duration human exploration missions of space. Therefore, NASA’s Langley Research Center and Johnson Space Center are building a low-voltage EB F3 system to test in a microgravity environment on board the KC-135A research aircraft. Considerations for KC-135A requirements and the effect of the microgravity environment on the hardware and process are built into the system design for the electron beam gun, vacuum, positioning, and wire feed.

Experiments are described to develop the processing envelope and understand the effects of microgravity on the geometry of the molten pool, the resulting build geometry, and solidification microstructure from the EB F3 process.464REFERENCES[1] Watson, K., D. Petersen, and R. Crockett, “Application of Solid Freeform Fabrication Technology to NASA Exploration Missions,” in Proceedings, 10th Solid Freeform Fabrication Symposium, Austin, TX: University of Texas at Austin, pp. 857 – 864, (1999).[2] Taminger, K. M. B., R. A. Hafley, and D. L. Dicus, “Solid Freeform Fabrication: An Enabling Technology for Future Space Missions,” in Proceedings, 2002 International Conference on Metal Powder Deposition for Rapid Manufacturing, San Antonio, TX: Metal Powder Industries Federation, pp. 51-60, (2002).[3] Taminger, K. M. B., and R. A.

51-60, (2002).[3] Taminger, K. M. B., and R. A. Hafley, “Characterization of 2219 Aluminum Produced by Electron Beam Freeform Fabrication,” in Proceedings, 13th Solid Freeform Fabrication Symposium, Austin, TX: University of Texas at Austin (2002).[4] Edwards, D. L., “Absorbed Dose Determination using Experimental and Analytical Predictions of X-Ray Spectra,” PhD Dissertation, August 1999, Department of Mechanical Engineering, Auburn University.[5] Hernandez, J., NASA-JSC Internal Memorandum, June 26, 2002.465.

VELOPMENT OF A QUALITY SYSTEM FOR POLYMER BASED- SELECTIVE LASER SINTERING PROCESSA. O. Stephen*, K. W Dalgarno*, J. Munguia* *School of Mechanical and Systems Engineering, Newcastle University, Newcastle, UKAbstractThe aim of this study is to develop a quality system for selective laser sintering, based on defining a minimum set of tests to qualify a build. MFI, impact and flexural tests were assessed, along with density, dimensional measurements and SEM. A benchmark part was designed for manufacture to track changes in key parameters from build to build, and tests on this validated against ISO standards. It is concluded that a combination of measures of flexural modulus, density and impact strength can be used for process monitoring and to infer the quality of a build in SLS process.1. IntroductionThe two methods of manufacturing that have dominated to date are subtractive (machining) or formative processes such as casting, injection moulding and forging.

IntroductionThe two methods of manufacturing that have dominated to date are subtractive (machining) or formative processes such as casting, injection moulding and forging. However, more recently the growth in three dimensional computer aided design software has enabled the production of product directly from 3D CAD models by material consolidation in layers without the need of tooling or jigs [1]. This has simplified 3D part production processes to 2D layering processes [2]. This new set of manufacturing methods is called additive manufacturing (AM). The building methods for additive manufacturing includes 3-D printing (3 DP), fused deposition modelling (FDM), laminated object manufacturing (LOM), selective laser sintering (SLS), selective laser melting (SLM) and stereolithography (SLA) [3-5].

The various processes are cost effective methods for the production of custom made parts based on customer requirements and are also good for low volumes with potential for medium and high volume production [6, 7].Parts made by these processes initially were for prototypes, concept verifications and analysis but technical improvements, better process control and the ability to use more materials resulted in a shift to rapid tooling and more recently to rapid manufacturing [8, 9].The SLS process uses a laser that sinters selectively a thin layer of powder spread over a platform. A computer directs a laser scanning mirrors over the powder layer, sintering and attaching a new layer of the part [4]. Each time a layer is finished, the platform is lowered and a new layer of powder is spread over the previously built layer.

Each time a layer is finished, the platform is lowered and a new layer of powder is spread over the previously built layer. This process is repeated sequentially until the part is completed and the sintered product is then separated from un-sintered powder after the cooling down stage.The common practice in Laser Sintering-based systems involves the blending of virgin powder with used powder to reduce cost and increase material utilisation. However part properties have been shown to later reduce as a result of repeated exposure to heat [10].This results in different chemical and structural properties and consequently lower physical and mechanical properties of the final part. A powder life study was carried out by Choren et al. [11] by increasing laser power relative to powder age, and it was observe that although increase in laser power increases most mechanical properties but higher powers and older powders chemically degrade the powder.Gornet et al.

[10] used an extrusion plastomer (a device that is normally used in plastic industries to measure the quality of resins and grade polymer) to assess powder properties with repeated use. It was observed that after several builds the melt flow rate (MFR) of polymer decreased with number of builds, an indication that the polymer is degrading. Low melt flow index was also attributed to improvement of elongation at break. Melt - flow rate (MFR) or melt volume - flow rate (MVR) is the rate of extrusion of a thermoplastic or resin through an orifice of standard dimensions under prescribed temperature and pressure [12].Various researchers have carried out work in improving the quality of fabricated part in addition to the aforementioned. Schmid and Levy [13] presented a generic model for the development of quality management system for additive manufacturing with selective laser sintering process.

Schmid and Levy [13] presented a generic model for the development of quality management system for additive manufacturing with selective laser sintering process. Real time melt pool analysis and control to achieve desired quality through the use of feedback control system in powder based SLS processing technology was proposed by Berumen et al. [14] for metal parts. Krauss et al.[15] also used thermography for monitoring of process parameter deviation in selective laser melting. An online quality control system for selective laser melting by the use of systems for monitoring powder layers deposing and real time melt process has also been developed [16]. However no correlation was made between the evolved material properties and the process parameters.The various research efforts point to the fact that a quality management system is very important for the future of additive manufacturing as the standardization activities are still at infancy level [13].

Overall the need of ensuring that AM processes deliver reliable and predictable properties cannot be overemphasised. Currently, the inability to guarantee material properties is holding back the adoption of AM technologies as industry does not have the confidence that manufactured parts will have the required mechanical properties for specific needs. This paper will therefore look at basic mechanical property variation with builds with the aim of developing a quality system for polymer based SLS systems.2. Testing Rationale2.1. Requirements for a QA testQuality assurance (QA) is defined as various organized procedures that ensure quality requirements for a product are satisfied. However, since it involves resources there is need of ensuring that the QA test will be simple, small, reliable, cheap and capable of being carried out quickly.2.2.

However, since it involves resources there is need of ensuring that the QA test will be simple, small, reliable, cheap and capable of being carried out quickly.2.2. Evaluation of Possible QA TestsA benchmark part was designed and fabricated to capture the key quality characteristics, after a review of various design. One benchmark sample design patterned after ISO designs for bending modulus and impact strength is involved in this study.Figure 1 shows the design for Benchmark part with dimensions all in millimetres.Benchmark (thickness of 4 mm) Figure 1 the geometrical design for Benchmark part3.

Experimental proceduresNylon SLS specimens were made by Peacocks Medical Group, Newcastle in Duraform PA 12 material, using a 3DSystems sProSD SLS machine, and with the processing conditions shown in Table 1.Table 1 SLS process parameters for parts supplied by Peacocks Medical GroupEquipment Laser power Outline laser power Fill scan spacing Laser scan strategy Layer thickness Scan speed3D Systems sPro60SD 12 W 6 W 0.15 mm Cross Fill Scanning 0.1 mm 5 m/s3.1. Test specimensFlexural specimens (Figure 2) were made in accordance with BS EN ISO 178:2010 [17] with the tests conducted using an Instron 4505 electromechanical system at a cross speed of 2 mm/min and were oriented in X, Y and Z directions as shown in Figure 2. Fifteen specimens with each five oriented in X, Y and Z directions were made for the flexural properties determination.Izod notch impact test specimens (Figure 3) were made also according to BS EN ISO180:2001 method A [18].

Tests were then carried out using a pendulum impact tester with pendulum capacity of 22 J. Fifteen specimens were made in a pack of five and oriented in X, Y and Z directions as shown in Figure 3 to reduce variability in thermal history due to locations in the build chamber by beams of thin section, which can be easily separated from each specimen after build. The build setup is also shown in Figure 2.(a)Z(b)YXFigure 2 (a) Five flexural specimen parts with dimensions (b) build setup with orientationsISO sample sizeFigure 3 Five impact specimen parts with dimensions3.2. Test Procedures for Benchmark Samples were clamped in a jig (fabricated using Rapid manufacturing (SLS) with EOS PA2200 material).

Test Procedures for Benchmark Samples were clamped in a jig (fabricated using Rapid manufacturing (SLS) with EOS PA2200 material). The testing procedures involved the placing of the benchmark sample in bottom block with a groove that fit the geometry of the sample and then placing of the top block on the sample such that the benchmark sample is sandwiched between the top and the bottom block of the jig. They are then clamped together in position with the aid of aluminium plate, two 100 mm bolts and wing nuts. Digital dial gauge is clamped vertically to sliding horizontal brackets whose position can be adjusted based on the length of the beam. The loads of 100g and 150g were hanged sequentially on the beam and readings of deflection for each load were then taken using Mitutoyo digital indicator. The data obtained was then used to compute the modulus for the beam.Figure 4 Positioning of the specimen during bending3.3.

The data obtained was then used to compute the modulus for the beam.Figure 4 Positioning of the specimen during bending3.3. Determination of density of specimensApproximate density of specimens was computed from measured volume obtained by measuring the dimensions (length, width and thickness) of the benchmark specimens using Mitutoyo digital calliper, which was then used for calculating the volume. The mass was measured with a precision balance (KERN PFB 200-3).4. Experimental results4.1.

Usually higher density results into superior strength due to reduction in porosities or voids that can serve as points of stress concentration. Also, there is greater possibility of crack initiation and failure for lower density sintered part due to presence of more un-sintered powder particles in the part [20, 21].From Figure 5 it can be observed that there is direct relationship between the density and the bending modulus for the benchmark samples. The correlation coefficient (Figure 6) is 0.932 with p – value of 0.021 which is less than the alpha level of 0.05 (above which the model can be said not to fit well to the data) [22]. The variation in densities may be due to a number of reasons some of which include: low powder bed temperature, energy density [19, 23], and interactions between them.Similarly, impact strengths (Figure 7) vary with densities and build with the exception of build five for the benchmark sample.

However, correlation between impact strengths and densities (Figure 8) was not statistically significant (p –value = 0.722) which may be due to surface flaws. Furthermore, molecular orientation and weight are also known to influence strength of polymer [24]. This may have been the reason for increase in impact strength in build five (with low MFI) in addition to increase in density. Also molecular weights has been suggested to be responsible for increase in elongation at break for laser sintered part [10, 25]. However, there is optimum range of molecular weight for laser sintering polymers as pointed out by Goodridge et al.[21] and Kruth et al.[26] beyond which the material properties will deteriorate. Thus the density of sintered parts, molecular weight and their interactions could be say to influence the mechanical properties of laser sintered part.5.2.

Thus the density of sintered parts, molecular weight and their interactions could be say to influence the mechanical properties of laser sintered part.5.2. Correlations of ISO impact strength and flexural modulus with that of BenchmarksamplesFrom Figure 9 the correlation coefficient between Benchmark and ISO impact samples is 0.96 with p –value of 0. Similarly, the correlation coefficient between the flexural modulus of Benchmark and ISO samples is 0.83 with the p – value of 0. Thus, correlation exists between the benchmark tests and ISO tests.6.

Thus, correlation exists between the benchmark tests and ISO tests.6. ConclusionFlexural modulus and impact strength were found to vary with density and the number of builds with a strong positive correlation observed between benchmark flexural modulus and density.Correlation between Benchmark specimens and ISO specimens in the same build wascarried out with positive correlation observed between them.In this study also, benchmark parts have been designed, manufactured and validated against that of ISO designed test specimens. The benchmark samples have demonstrated ability to capture the key quality characteristics of strength, stiffness and dimensional accuracy (the minimum resolution that the benchmark can capture is 1.2 mm) with number of builds. This therefore, opens a prospect for its usage in process monitoring and control. Further work is on-going in the implementation of the quality system with another materialand in partnership with an industrial partner.

Further work is on-going in the implementation of the quality system with another materialand in partnership with an industrial partner. Within a quality system, one benchmark specimen is recommended to be added to each build and placed about the same region in a build volume. The aforementioned quality characteristics can then be used independently or jointly depending on the desired monitor/control level that is required by the user and the available test resources. The data could also be used in statistical process control. Therefore, on the basis of the data from this study a proposed novel and cost effective quality system will therefore consist of qualification of input materials by the use of MFR and qualification of a build by the use of benchmark samples (density, stiffness and dimensional accuracy).

velopment of a Selective Laser Reaction Sintering Workstation B.R. Birmingham, J.V. Tompkins, G. Zong, and H.L. MarcusCenter for Materials Science and Engineering The University of Texas at Austin Austin, Texas 78712AbstractThe purpose of this paper is to describe the design and operationof a Selective Laser Reaction Sintering workstation developed at The University of Texas. The workstation allows the study of solid freeform fabrication of reaction sintered materials on a research scale. The mechanical and control systems of the workstation are detailed, and Selective Laser Reaction Sintering as a technique is discussed including example material systems that are currently under study.IntroductionSolid freeform fabrication (SFF) is any manufacturing process that produces athree dimensional part without the use of standard or part specific tooling. One form of SFF, Selective Laser Sintering (SLS), uses a laser to sinter selected areas of a powder bed.

One form of SFF, Selective Laser Sintering (SLS), uses a laser to sinter selected areas of a powder bed. Additional layers of powder are spread on top of the frrst, and specific areas of each layer are scanned and sintered by the laser before the next layer is spread. In this way, a three dimensional sintered part is built up layer by layer. At the end of the process, the manufactured shape may be removed from the surrounding loose powder. Parts have been made from various polymers, metals and ceramics using this technique [1,2]. Selective Laser Reaction Sintering (SLRS) combines SLS with a simultaneous gas/powder reaction process known as reaction sintering or reaction bonding [3]. This simultaneous reaction typically involves the decomposition of a gas to solid or a gas interacting with a liquid or solid to produce another solid.

This simultaneous reaction typically involves the decomposition of a gas to solid or a gas interacting with a liquid or solid to produce another solid. The SLRS process has the potential of producing parts made of monolithic materials and composites that are difficult or impossible to sinter using SLS as well as producing microstructures that are unobtainable using standard sintering techniques.EquipmentTo be successful a SLRS system must accomplish three basic functions. These functions are 1) supply controlled laser power to a selected area, 2) provide multiple layers of a powder bed having proper thickness and density and 3) provide the141appropriate gas environments including vacuum. An overall schematic of the SLRS system is seen in figure 1.C02 Lo.serHe-Ne Lo.serEnvironMento.lZnSe 'WindowRoller Tro. verse,--_---'----'-__~ Go.s InletsServo Driven XYZ Plo.

An overall schematic of the SLRS system is seen in figure 1.C02 Lo.serHe-Ne Lo.serEnvironMento.lZnSe 'WindowRoller Tro. verse,--_---'----'-__~ Go.s InletsServo Driven XYZ Plo. t f orMPowder Feecl o.ndelective La-ser Reactionsta- tionFigure 1.With the exception of the powder delivery mechanism, the basic arrangement ofthis system has been previously described in the literature [4]. The laser beam from a 25 watt C02 laser is directed onto a substrate that is located in an environmentally controlled chamber. A motor driven X-Y table moves the chamber/substrate under the stationary beam effectively "scanning" the beam. Both laser power and X-Y table controls have been upgraded for this workstation. Laser power fluctuations have been reduced to less than. 5 % of setting by using a pulse width modulation controller developed at the University [5]. Power level can be adjusted manually or by computer.

5 % of setting by using a pulse width modulation controller developed at the University [5]. Power level can be adjusted manually or by computer. Positional accuracy of the table/beam is now better than 20ILm and repeatability better than 5ILm at scan speeds to 3mm/sec using a computer controlled dc servo system. Environmental control is achieved by locating the powder deliverymechanism/substrate inside a vacuum chamber. The chamber has three independent gas inputs with the emphasis on flexibility due to the research nature of the system. The gas inputs have allowed for the introduction of a variety of gases including N2, N2/H2, NH3' Ar, H2, 02, CH4, and C2H2. The chamber is currently set up for static gas environments, but minor modifications could easily make it a flowing gas. system if desirable. Mechanical pumping initially provides 10-3 Torr vacuum to minimize gas impurity content. Mechanical and resistance vacuum gauges monitor system pressure.

system if desirable. Mechanical pumping initially provides 10-3 Torr vacuum to minimize gas impurity content. Mechanical and resistance vacuum gauges monitor system pressure. Power and thermocouple feedthroughs provide the ability to heat substrates and monitor and control temperature using a PC based data acquisition system. A gas sampling port has also been included to enable connection to a gas analyzer system(RGA). This will be used to gather information about reactants and by-products during the reaction sintering process. Mechanical feedthroughs currently enable manual operation of the powder delivery mechanism. The mechanism will be automated in the future.148The powder delivery system is similar in concept to one previously reported [6],but considerable miniaturization was required to fit the entire mechanism inside the available five inch diameter by 5 inch in height vacuum chamber. A schematic and photograph can be seen in figures 2 and 3.

A schematic and photograph can be seen in figures 2 and 3. Two rotation feedthroughs marked Powder Feed and Powder Accept are coupled to worm/gear/leadscrew arrangements that ultimately drive the powder feed and powder accept pistons up and down in their respective cylinders. The rotation feedthrough marked Roller Traverse drives a spur gear/leadscrew arrangement that causes the roller to traverse across the top of the stage.'0Powcler Spreacling RollerPowcler Accept PistonPowder Feecl PistonPow·der Delivery SysteMFigure 2.During a standard production cycle the system starts with the traversing roller to the far left, the powder feed piston would be down and its cylinder full of powder, and the powder accept piston would be up. Powder is made available to the roller by raising the powder feed piston an incremental amount. The roller is then traversed across the149stage and then back, spreading the powder into a thin layer.

The roller is then traversed across the149stage and then back, spreading the powder into a thin layer. As the roller traverses, a rack and pinion causes the roller to spin so that the roller surface in contact with the powder is actually moving in a direction opposite to traverse direction. This is done for improved powder spreading. After the new powder layer is spread, a laser scan is performed, sintering specific areas of the powder layer above the powder accept piston. When the scan is complete, the powder accept piston is lowered some specified layer thickness. The powder feed piston is again raised, providing another incremental amount of powder. The roller is brought across the stage, spreading a fresh layer of powder over the powder accept piston and the previously sintered layer. The laser is scanned again resulting in a second sintered layer. The powder accept piston is lowered and the process is repeated until the part has been built up layer by layer.

The laser is scanned again resulting in a second sintered layer. The powder accept piston is lowered and the process is repeated until the part has been built up layer by layer. Maximum part size from this mechanism is roughly a 1.5cm by 1.5cm by 1.5cm cube. The gearing of the mechanism enables layer thickness control to within a few microns. Working on this scale is advantageous when using difficult to make, hazardous or expensive materials because of the small amounts of precursor required to load the system. When the materials systems have been proven on this scale they can easily be adapted to a larger system.ApplicationsLaser reaction sintering can be used in a variety of ways to produce a variety ofresults. The following discussion is by no means all inclusive, but is meant to give examples demonstrating some of the interesting aspects of this technique.Standard pressureless sintering of oxides, nitrides, and carbides is difficult.

However, formation of an oxide, nitride or carbide may be achievable by reaction sintering in the corresponding oxygen, nitrogen, or carbon rich atmosphere. One example is the reported success of sintering an aluminum/alumina mixture in the presence of oxygen to form an alumina pre-form [7]. These pre-forms will be infiltrated to create a metal-ceramic or ceramic-ceramic composite part.Composite structures can be produced directly by partial conversion of powder by reaction. A multi-layer Cu-TiN part was produced by this SLRS system using a Cu tOTi alloy as base material. Reaction sintering in a nitrogen atmosphere caused the Ti to migrate to the surface of the alloy particles where it reacted with nitrogen to convert to TiN. The result was a copper matrix with a sub micron TiN layer on each of the original powder particles. TiN presence was confmned by x-ray diffraction and the dispersion of Ti rich areas(TiN) was mapped using EDS.

TiN presence was confmned by x-ray diffraction and the dispersion of Ti rich areas(TiN) was mapped using EDS. Figure 4 shows a side view of a 6mm by 6mm by 3mm thick Cu-TiN rectangular solid made using SLRS. Note that the part is composed of 15 layers. The delamination observed can be eliminated with proper substrate pre-heating. The microstructure of the part can be seen in Figure 5, which contains a backscattering electron micrograph and a Ti element map of the same region. The element map clearly shows the Ti migration to the surface of the alloy particle where it formed a nitride coating on the particle. It is believed that proper control of powder size and operating parameters can result in very fine microstructures unobtainable by standard sintering techniques.150A third technique involves infiltration.

One system currently from an organometallic pyrolitically produced SiC infiltrates and multi-layer a single layer of through the rectangular the layers parameters J.1I.J........uuJ.up;,simultaneous vapor deposition andpyrolitic formation of SiCa SiC base powder. TheOO'WQ~~r and binds it together. Single "'1101111"&1> 6 shows the side view of Note the density variationraSJrllOln..:n..lIU:«.LIl..., by Imm thick SiC16p,m SiC. Bonding between 1 rn1~1"n'U&l>.rl by optimization of processand layer thickness.Figure 7.ConclusionsA Selective Laser Reaction Sintering workstation isand operational. Itis capable of producing multi-layer solid freeform parts in an environmentally controlled chamber. This will enable the study of laser sintering combined with gas of producing phase/powder that this combination Preliminary materials and structures unobtainable by standard sintering studies are positive.It is152AcknowledgmentsThanks go to Dr.

Uday Lakshminarayan for his early work on this program, EdBarth, and Dr. Michael Schmerling for their microscopy assistance and to the Texas Advanced Technologies Program (ATP) project 323 and the Office of Naval Research grant NOOOI4-90-J-II64(Cu-Ti) and NOOOI4-92-J-1514(SiC gas/ SiC solid) for funding this effort.References1Harris L. Marcus, Joseph J. Beaman, Joel W. Barlow, and David L. Bourell, "From Computer to Component in 15 Minutes: The Integrated Manufacture of Three-Dimensional Objects," JOM,42[4], 1990,pp. 8-10.2"Rapid Prototyping Systems," Mechanical Eniineering, edited by Steven Ashley, April 1991, pg 41.3M.E. Washburn, W.S. Coblenz, "Reaction-Formed Ceramics," Ceramic Bulletin, Vol 67, 2, 1988, pp. 356-363.4Guisheng Zong, R. Carnes, H. G. Wheat, and H.L. Marcus, "Solid Freeform Fabrication by Selective Area Laser Deposition," Proceedings of the Solid Freeform Fabrication Symposium, Austin, Texas, August 6-8, 1990, pp. 83 90.5J. V.

Wheat, and H.L. Marcus, "Solid Freeform Fabrication by Selective Area Laser Deposition," Proceedings of the Solid Freeform Fabrication Symposium, Austin, Texas, August 6-8, 1990, pp. 83 90.5J. V. Tompkins, "Computer Control of Laser Intensity for Selected Area Laser Deposition," Masters Thesis, University of Texas at Austin, Austin, Texas, 19926C. Deckard and J.J. Beaman; pp. 623-629 in Advances in Manufacture, Integration and Processes, Proceedings of the 15th Conference on Production Research and Technology, Berkeley, California, January 9-13, 1989. NSF Design, Manufacturing and Computer Integrated Engineering Division, Washington D.C., 1989.7P.K. Subramanian, G. Zong, and H.L. Marcus, "Selective Laser Sintering and Reaction Sintering for the Fabrication of Ceramic Composites," Proceedings of The Solid Freeform Fabrication Symposium, Austin, Texas, August, 1992.153.

E DEVELOPMENT OF A SLS® COMPOSITE MATERIALPaul Forderhase Kevin McAlea Richard Booth DTM Corporation Austin, TX 78754ABSTRACTThe development of a commercial SLS nylon-based composite material (LNC 7000) is described. Nylon composite candidate systems with different volume fractions of a number of glass fiber and glass bead reinforcements were screened. It was found that fully dense SLS parts with excellent mechanical properties could be made from a number of reinforced nylon materials. An optimized material containing 29 volume percent 35 Ilm diameter glass beads was selected based on the processing behavior and mechanical properties of the candidate systems. The performance of this optimized material is In addition, complementary aspects of the composite nylon and unreinforced described.

The performance of this optimized material is In addition, complementary aspects of the composite nylon and unreinforced described. nylon materials (LN 4010 and LNF 5000) are discussed.1.INTRODUCTIONComposite materials are widely used in the polymer industry, particularly where increased stiffness and strength are required. There are many types of polymer composites. The reinforcement phase can be in the form of continuous roving, a woven fabric, short fibers, irregular particles, or regular spheres. Expensive, high performance reinforcements such as carbon/graphite are available, as are inexpensive, lower performance reinforcements such as glass. The polymer binder phase can be thermoset or thermoplastic, depending on the type of process used to make the part and the desired properties.

The polymer binder phase can be thermoset or thermoplastic, depending on the type of process used to make the part and the desired properties. Injection molded composites, which typically consist of a ductile, thermoplastic binder reinforced with short fibers or a particulate, are a particularly important type of polymer composite.Selective Laser Sintering (SLS) is now widely used to produce fully-dense functional prototypes from nylon (LN 4010 and LNF 5000). Nylon parts produced with the DTMTM SLS process have properties representative of some common injection molded engineering thermoplastic parts (1). These parts are produced directly from CAD data without support structures or post-curing (2).Composite material parts have been produced with SLS technology at the University Texas at Austin and at DTM Corporation.

These composite parts have typically been made from polymer coated metal and ceramic materials in which the polymer binder is less287than 20 volume percent of the part. These porous SLS parts are typically post-processed at high temperatures to obtain the final desired properties (3 - 6).In this paper, the development of a SLS composite material (LNC 7000) based on nylon is described. The goal was to produce full density parts with properties more representative of injection molded thermoplastic composites. In contrast to previous SLS composite work, the polymer binder in this material is greater than 70 volume percent of the part. The experimental screening and optimization, and work performed to develop the commercial system are described, as are the performance and processing advantages of this material.2.

The experimental screening and optimization, and work performed to develop the commercial system are described, as are the performance and processing advantages of this material.2. EXPERIMENTAL2.1 Screening Nylon composite candidate systems with different volume fractions of The screening various glass fiber and glass bead reinforcements were screened. since only experiments were limited to composites containing fibers or beads, discontinuous reinforcements are compatible with current SLS powder feeding. In addition, only glass reinforcements were examined because of their availability and cost effectiveness. Parts from the screening experiments were evaluated based on mechanical properties, ease of processing, breakout, and post-finishing.

Parts from the screening experiments were evaluated based on mechanical properties, ease of processing, breakout, and post-finishing. The screening experiments were followed by an optimization of the SLS nylon composite system based on the above criteria.2.2 Fiber Reinforcements Fibers are a relatively efficient reinforcement geometry, typically increasing the stiffness, strength, and heat deflection temperature of parts relative to the unreinforced polymer. Two different types of glass fibers were screened for use as a SLS nylon reinforcement. The first fiber had a mean length of 85 flm, and a mean diameter of 10 flm, while the second had a mean length of 70 flm, and a mean diameter of 16 flm. The composite material based on the smaller diameter fiber exhibited poor bulk flow properties. One of the material requirements for SLS processing is that the powder flows freely, even at elevated temperatures, since powder flow is required to form each new layer.

One of the material requirements for SLS processing is that the powder flows freely, even at elevated temperatures, since powder flow is required to form each new layer. It is speculated that the poor flow of this system was due to the aspect ratio (8.5) of the fiber which results in significant fiber-fiber contacts and considerable fiber entanglement. The powder flow properties of the composite prepared with the larger diameter fibers were not adversely affected. The aspect ratio of this fiber material is 4.4. Composite parts with this fiber were fabricated with reinforcement levels of 9.6, 15, and 22 volume percent.The fiber reinforced parts all showed greater modulus and strength compared to unreinforced nylon. SLS composite parts reinforced with 15 volume percent fiber showed an increase in tensile modulus of approximately 50%, and an increase in tensile strength of about 25%. The fiber composite parts were also found to exhibit complex anisotropy.

The fiber composite parts were also found to exhibit complex anisotropy. Since the mean fiber length was 70 Jlm and the typical layer thickness approximately 100 flm, some degree of in-plane anisotropy would be expected. However, it was also found that stiffness and strength in the x direction (the direction of roller travel) were288significantly higher than in the y direction (transverse to the direction of roller travel) as shown in Figure 1. This effect is consistent with fiber alignment in the direction of the roller travel during leveling.

This effect is consistent with fiber alignment in the direction of the roller travel during leveling. The data in Figure 1 were generated by testing the tensile properties of specimens reinforced with 15 volume percent glass fibers built both in the x and y directions.I""""i'N Nr---;-'lJ) 'I:j-0-i'I:j- 'I:j-(U CL ~C'0 'I:j-N 'I:j-I(J) :J "'5 u 0 ~0 N~~~I~~xyxyOrientationOri entati onFigure 1: Anisotropy in Short Fiber Nylon CompositeThis work clearly demonstrated that it is possible to produce SLS polymer composite parts with a short fiber reinforcement, with certain restrictions on the fiber geometry. Since the scale/offset compensation software allows different scale values to be applied in the x, y, and z directions, it should also be possible to build accurate parts with such a material.

Since the scale/offset compensation software allows different scale values to be applied in the x, y, and z directions, it should also be possible to build accurate parts with such a material. It was decided, however, to pursue isotropic systems which would offer many of the advantages of a fiber reinforced material without the complications of anisotropic shrinkage and mechanical properties.2.3 Glass Bead Reinforcements Glass beads are a less efficient reinforcement than glass fibers, but parts made with spheres exhibit isotropic properties. Polymer parts reinforced with glass beads typically have a much higher modulus than unreinforced parts, but show only a small improvement in strength. In initial screening work, it was found that fully dense, isotropic parts with improved modulus could be produced with a number of glass bead materials.

In initial screening work, it was found that fully dense, isotropic parts with improved modulus could be produced with a number of glass bead materials. Further optimization work focused on materials made with a range of reinforcement levels and glass bead diameters.2892.3.1 Effect of Glass Bead Reinforcement Level Glass bead reinforced composites with reinforcement levels over the range 9 to 39 volume percent were examined. The modulus of a composite material will generally increase as the volume fraction of reinforcement increases until a packing limit is reached. The relationship between the modulus of a particulate reinforced composite and the reinforcement volume fraction can be modeled by (7):PPm=(l+~l1vf)(1-11 v f)(1)11 =( PI PmIJ(2)( PI Pm+SJWhere p is the modulus of the composite, pm is the modulus of the binder, and Pf is the modulus of the reinforcement.

The factors ~ and 11 are parameters which depend on the boundary condition between the matrix and reinforcement as well as the moduli of the matrix and reinforcement. The assumed value of ~ for this type of reinforcement in this regime of reinforcement volume fraction is 2. The experimental and theoretical values of tensile modulus are plotted as a function of volume percent reinforcement in Figure 2. The experimental data are for composites made with a mean glass bead diameter of 35 Ilm, although the modulus is not a function of the bead size.2903.53(i' a.. ~ fh 2.5 ::::I :i "0 0 ::it .! '(ji c {!!.21.5/'"~/:V~/ ~411+- Theoretical ____ Experimental1 0.050.10.150.20.250.30.35Volume Fraction of Reinforcement (%)Figure 2: Theoretical and Experimental Values of Tensile ModulusThe experimental values were found to be in reasonable agreement with those predicted by the model up to a reinforcement level of approximately 29 volume percent.

Up to this level of reinforcement, the composite modulus increases monotonically with increasing volume fraction reinforcement as expected. Above a reinforcement level of approximately 30 volume percent, parts were found to exhibit considerable porosity and the modulus/reinforcement level relationship predicted by Equation 1 was found to be no longer valid.2.3.2 Effect of Glass Bead Diameter Composite materials with mean glass bead diameters ranging from 4 to 114 Jlm were studied. Materials made with bead diameters of 4 and 11 Jlm were found to exhibit poor bulk flow at high temperatures, presumably due to the higher interparticle friction found in extremely fine powders. As a result, only parts with bead diameters of 35,49, and 114 Jlm were made and evaluated.relationships between glass bead diameter and tensile strength and glass bead diameter and ultimate elongation are shown in Figure 3.

These samples all contained 29 volume percent reinforcement.291~r--;-oco 0 0l{)""'"0""'",~~i~iiic 0 ~ 0)c 0 [jJ ~ E E =:J<0 ~ 0""'" ~ 0l{) <:0~r--r-'l~ I.-1.....J<"l 0 0~11449351144935Average Bead Diameter (microns)Average Bead Diameter (microns)Figure 3: Comparison of Mechanical Properties by Reinforcement Bead Average DiameterThe tensile strength and ultimate elongation both increase with decreasing bead diameter. This is not surprising, since smaller size reinforcing particles generally provide a more homogeneous stress distribution and improved ultimate properties in composites (8).A typical tensile fracture surface for a part made with 29 volume percent, 35 J.Lm bead diameter glass beads is shown in Figure 4.292Figure 4: Fracture Surface of Nylon CompositeWhile there is some evidence of mixed-mode failure, the dominant failure mode appears to be debonding of the glass beads from the nylon polymer.

It is speculated that the absence of consolidation pressure during the SLS process is primarily responsible for the limited adhesion. The limited adhesion probably contributes to the relatively high ultimate elongation of these composites, since failure occurs through gradual damage accumulation rather than catastrophic crack initiation and propagation.3.1 DTM Nylon Composite: LNC 7000 Based on the data presented above, a nylon composite material reinforced with 29 volume percent 35 J!m glass beads was chosen for commercialization. This material offers high modulus with good strength and ultimate elongation. Parts produced from this composite are shown in Figure 5. It was also found that the nylon composite system offers SLS processing advantages.

Parts produced from this composite are shown in Figure 5. It was also found that the nylon composite system offers SLS processing advantages. In particular, the nylon composite material offers a broad processing window, reduced part warpage, easy part breakout, and easy finishing.3.2 Nylon Composite Mechanical Properties The properties of the commercial nylon composite material (LNC 7000) are compared to those for unreinforced nylon (LN 4010 and LNF 5000) in Table 1.293Figure 5: LNC 7000 Nylon Composite PartsTable 1: Comparison of SLS Nylon Composite (LNC 7000) To Fine SLS Nylon (LNF 5000)Property Tensile Stren2th Tensile Modulus Tensile Elon2ation Flexural Modulus Notched Izod Unnotched Izod Deflection Temp.

Under Load (1.8 Mpa)Unit MPa MPa % MPa JIm JIm °CTest Method L Fine Nylon~Composite 36 ASTMD638 1400 ASTMD638 32 ASTMD638 870 ASTMD790 70 ASTMD256 1370 ASTMD256 44 ASTMD64849 2828 6 4330 68 443 134The tensile modulus of the composite is approximately twice that of the unreinforced nylon, while the flexural modulus of the composite is approximately five times greater than that of the unreinforced nylon. The tensile strength for the composite is approximately294I30% higher than that of unreinforced nylon, while the heat deflection temperature of the composite is 90° C higher than unreinforced nylon. As expected, the tensile elongation and notched Izod values for the composite are lower than those of the unreinforced nylon due to the presence of the brittle glass phase.

As expected, the tensile elongation and notched Izod values for the composite are lower than those of the unreinforced nylon due to the presence of the brittle glass phase. The two nylon systems are complimentary; the unreinforced material offers outstanding ductility with good strength and heat resistance, while the composite material offers outstanding stiffness with even higher strength and heat resistance.3.3 Processing Window The nylon composite material has a 3 to 4°C processing "window" (defined as the part bed temperature range over which a part can be successfully built) compared to only 1° C for the equivalent unreinforced nylon. This greater process latitude is probably due to the higher feed temperature (10° C) and lower part bed temperature (2 to 4°C) setpoints available with the nylon composite. As the difference between the feed and part bed temperature is reduced, there is less chance of "in-build" part warpage which is a primary SLS build failure mode.

As the difference between the feed and part bed temperature is reduced, there is less chance of "in-build" part warpage which is a primary SLS build failure mode. Within this 3 - 4° C temperature window, there is a striking interaction between part bed temperature and laser power which allows the user to adjust part density and detail resolution to the desired value. At the low end of the temperature window, higher laser power can be used to attain the same density that lower laser power will produce at the higher temperature. This behavior is illustrated in Figure 6.Figure 6: Laser Power Part Bed Temperature InteractionThe two columns of parts in Figure 6 correspond to two part bed temperature setpoints (194° C, and 195° C, from left to right) while the rows correspond to three laser power setpoints (0.75, 1, and 1.5 Watts, from bottom to top).3.4 Part Breakout Nylon composite parts are generally easier to remove from the part cake (breakout) than unreinforced parts.

This relatively easy breakout can probably be295attributed to reduced "growth" (unwanted powder sintering at part boundaries) which is the result of the lower part bed temperature and reduced laser power (approximately 2 watts vs. 4 watts) used with the nylon composite. In addition, the reinforced material probably has less intrinsic potential for growth due to the presence of the glass phase which cannot sinter at the nylon composite part bed temperature.3.5 Part Warpage Parts made from nylon composite typically show less "post-build" warpage (approximately one half) than the corresponding parts made from unreinforced nylon. The composite parts shrink less than unreinforced parts (3% vs.

The composite parts shrink less than unreinforced parts (3% vs. 4%) which probably reduces their tendency to develop stresses and warpage as they are cooled the build temperature to the breakout temperature.3.6 Post-Finishing Nylon composite parts can typically be finished in approximately one half the time it takes to finish corresponding unreinforced nylon parts. The presence of the hard, brittle glass phase makes it easier to sand the composite parts. A typical nylon Jlm RMS in 5 composite part can be finished to an average surface roughness of hours.4. CONCLUSIONSFully dense SLS parts with excellent mechanical properties can be made from nylon and a number of fiber and bead reinforcements.A commercial nylon composite material (LNC 7000) with 29 volume percent 35 flm glass bead reinforcement is now available.

This material offers increased stiffness, strength, and heat resistance compared to unreinforced nylon as well as easy processing and post-finishing.The family of SLS nylon materials - unreinforced nylon (LN 4010 and LNF 5000) and composite nylon (LNC 7000) - offers a range of properties representative of many injection molded engineering plastics.2965. REFERENCES1. P. Forderhase, M. Ganninger, and K. McAlea Proceedings of the 4th European Conference on Rapid Prototyping and Manufacturing The University of Nottingham, Nottingham UK, 1995 2. P. Forderhase, K. McAlea, M. Michalewics, M. Ganninger, and K. Firestone Proceedings of the Solid Freeform Fabrication Symposium The University of Texas at Austin, Austin, TX, 1994, p. 102 3. I. Lee, A. Manthiram, and H. Marcus, Proceedings of the Solid Freeform Fabrication Symposium, The University of Texas at Austin, Austin, TX, 1994, p. 339 4. N. Vail, J. Barlow, and H.

Using in-situ monitoring, the identification of defects, as well as the geometric and dimensional measurement of the layers throughout the build allows for greater quality control, as well as a reduction in the requirement for ex-situ measurement. A standalone monitoring system for the EOS M280 is presented in this research, allowing for the build process to be monitored layer-by-layer. The system images the build area after powder deposition and after laser exposure allowing for the identification of inefficiencies in both the powder deposition and the laser exposure. The system has proven to be capable to identify in build defects and work is ongoing to develop an automated program to identify these defects and notify the operator in real time.

The system has proven to be capable to identify in build defects and work is ongoing to develop an automated program to identify these defects and notify the operator in real time. Keywords: Metal additive manufacturing, Powder bed fusion, in-situ process monitoring, defect detectionIntroductionAdditive manufacturing (AM) allows for the manufacture of high-value complex components and as such has gained popularity in a range of industries. Due to this, research into the AM process has received increased attention. The AM process is a layer-by-layer process allowing for the production of high value, complex components in industries such as aerospace, automotive and medical device [1], [2]. The process presented in this research is a powder bed fusion (PBF) process called Direct Metal Laser Sintering (DMLS). The process selectively melts a powder bed, layer-by-layer to build a final part. The powder layer is deposited on the build platform by the re-coater blade.

The process selectively melts a powder bed, layer-by-layer to build a final part. The powder layer is deposited on the build platform by the re-coater blade. The laser then melts the powder layer and a new layer of powder is deposited, as depicted in Figure 1. The process repeats layer-by-layer until the part is complete .Re-coater BladeGalvo-MirrorLaser Beam..LaserPartMetal PowderBuild PlatformOverflowM il CiFigure 1: DMLS process, showing the powder dispenser, build platform, laser system and AM part1390The ability to manufacture components while monitoring their quality throughout the process is a requirement for many of the highly regulated industries such as medical device, aerospace and automotive. This is pertinent to the quality assurance of the final AM part. This area ofresearch has been identified as one requiring significant attention [3].

This is pertinent to the quality assurance of the final AM part. This area ofresearch has been identified as one requiring significant attention [3]. The use of in-situ monitoring allows for in-process defects to be identified and quantified during the build process eliminating costly post-process testing such as computed tomography (CT) [4]. Quantifying the defect will allow for a decision to be made on the impact of the defect on the final integrity of the part. The application of in-situ monitoring can also be used to observe the consistency of the powder deposition process in DMLS. Consistent and repeatable powder deposition is vital for the successful outcome of the powder bed fusion (PBF) process as it leads to consistent melting during laser exposure [5] [6].The early work in the in-situ inspection of the powder bed fusion process utilized an in-line camera system [7].

Previous monitoring systems for the PBF process have utilized the application of some form of a digital camera mounted to the roof of the build chamber pointing at the build platform. This method is known as build platform monitoring which enables the build plate and distributed powder layer to be monitored throughout the build process. The camera images the build platform and the resulting powder bed allowing for defects and anomalies in the powder deposition and the laser scanning to be identified. Utilization of a high resolution digital single-lens reflex (DSLR) camera mounted within the chamber of a PBF machine enables images of the deposited powder bed to be captured, as well as the subsequent laser exposure. Abdelrahmn et al. [8] used this method to detect anomalies in the powder bed after re-coating and laser scanning. An algorithm was applied to the captured images in order to detect anomalies between layers. Land et al.

An algorithm was applied to the captured images in order to detect anomalies between layers. Land et al. [9] used a similar method; however, this paper proposed the use of multiple cameras in order to image the build platform from multiple angles. This multi-camera approach allows for the parts being produced to be captured at different angles permitting their features to be observed.Jacobsmtihlen et al. [1 O] used a charge coupled device (CCD) camera mounted on the outside of the chamber door of an EOS M270. Images were captured using various light sources after powder recoating. Areas which appear shiny in the images are those in which there is a lack of powder deposited. These areas were highlighted using an image analysis tool and the operator was notified.

Areas which appear shiny in the images are those in which there is a lack of powder deposited. These areas were highlighted using an image analysis tool and the operator was notified. This method illustrated that the setup can identify areas where there is insufficient powder deposited [10].Other in-situ monitoring methods have been applied to the PBF process which utilizes the use of thermal imaging to monitor process signatures. A process signature is a dynamic characteristic of the PBF process, for example the size, shape or temperature of the melt pool during laser interaction. These signatures can significantly affect the final quality of the parts produced [3]. Monitoring of the process signatures has been previously investigated in [11 ], [12] and [13]. Thermal imaging allows the process signature to be monitored and used to control the process parameters such as laser speed and power.

Thermal imaging allows the process signature to be monitored and used to control the process parameters such as laser speed and power. The ability to adjust the process parameters, laser speed and power, enables the process signature to be controlled [14].Detects in the PBF processes lead to the weakening of the mechanical properties which are critical in some of the industrial applications using metal AM processes [6]. The origin of the defects can be caused by the following [15]:a. Powder feedstock or chosen processing parameters. b. Ineffective build planning such as part geometry, support design and part orientation. c. Uncalibrated or damaged equipment.The objective of an in-situ monitoring system is to identify defects that originate from the list shown above. Some of the common defects in the PBF process have been previously classified into the following categories, defined by the way the defect affects the quality of the final manufactured part [6]:1.

Some of the common defects in the PBF process have been previously classified into the following categories, defined by the way the defect affects the quality of the final manufactured part [6]:1. Geometry and Dimension - Geometric and dimensional inaccuracy. 2. Surface Quality - Surface roughness, morphology and deformation.13913. Microstructure -Anisotropy, heterogeneity and poor density. 4. Mechanical Properties - Fracture, cracking, porosity and lack of interlayer bonding.These defects originate from the reasons outlined above. For the PBF process the resulting part quality relies on a consistent and uniform deposition of powder across the build plate. This ensures that the laser interaction is repeatable for each layer of the build.

This ensures that the laser interaction is repeatable for each layer of the build. Variations in the uniformity of the powder layer can lead to re-melting of previous layers or lack of fusion to the previous layers.This research will utilize the methods of line segment and anomaly detection to identify defects in the build platform. Line segment detection is a machine learning technique that identifies locally straight contours known as line segments in digital grey-scale images. Anomaly detection is the identification of patterns in data that do not conform to typical behavior. Failure to detect anomalies can cause loss of revenue and business as well as reputation. In smart manufacturing, data is gathered using sensors that monitor processing conditions. Detecting anomalies as they occur in production can indicate tool wear [16] and tear and whether a product is deviating from specification, thus, reducing scrap rates and lowering energy consumption [17].

Other applications of anomaly detection are ubiquitous in our data rich world, for example, fraud detection [18], telecommunications [19] , bio-surveillance [20] and infection control [21].Research in anomaly detection abounds in the scientific literature. Reviews and/or experimental studies are presented in [22], [23] and [24] from the statistical, machine learning and computer science perspectives, respectively. Gupta [25] provides a thorough review, comparing classification based, clustering based, nearest neighbor based, statistical and information theoretic approaches for anomaly detection using examples from the diverse fields of fraud detection, image processing, sensor networks and textual anomaly detection.Kandanaarachchi and Mu [26] discussed the algorithm selection problem for unsupervised anomaly detection. Alternatively, a combination of techniques may be applied.

Alternatively, a combination of techniques may be applied. Talagala et al, [27] provide a useful framework to optimize anomaly detection performance in high frequency time-series data. The framework is a 10-step process to aid the identification of end-user needs, prioritizing anomaly types and selecting suitable methods of anomaly detection, assessing and comparing their performance and making recommendations. An example of the use of this framework in water-quality data is given in [28].In this research a standalone in-situ monitoring system for defect detection in the PBF process is presented. The monitoring method presented is a standalone optical monitoring system for the PBF process allowing the powder deposition and laser exposure to be monitored. Despite the rapid developments and improvements in the PBF processes the range of defects are currently limiting the repeatability of the process [6].

Despite the rapid developments and improvements in the PBF processes the range of defects are currently limiting the repeatability of the process [6]. In-situ monitoring of the process has been identified as an area in need of further research for the additive manufacturing process [3].Development of the Monitoring SystemThe aim of the development of the monitoring system presented in this research is to provide a standalone platform allowing for the build plate to be monitored layer-by-layer throughout the PBF process of Direct Metal Laser Sintering (DMLS). The method presented in this paper allows for the identification of defects and inefficiencies in the powder deposition stage.

The method presented in this paper allows for the identification of defects and inefficiencies in the powder deposition stage. These defects may include but are not limited to:(cid:120) (cid:120)Lack of powder deposition (short-feeding) Raised edges or lifting partsThe system has been installed in the chamber of the EOS M280 DMLS machine in the South Eastern Applied Materials (SEAM) Research Centre in Waterford Institute of Technology. As the monitoring system is standalone from the operation of the M280, there are some constraints in the selection of the required image acquisition equipment. The installation of the selected system is presented in the following section.1392Installed SystemThe chosen system consists of a 21 MP complementary metal oxide semiconductor (CMOS) monochrome camera with a 16 mm lens attached. This is mounted in the chamber above the build platform, as shown in Figure 2, to capture images of the build platform during the AM process.

This is mounted in the chamber above the build platform, as shown in Figure 2, to capture images of the build platform during the AM process. The camera is triggered using two laser trigger sensors mounted behind the machine panels, as shown in Figure 2. These sensors utilize the motion and position of the re-coater blade to trigger the camera and capture the image. The images are then collected by the camera controller unit and stored to an external hard drive.-------Figure 2: In-situ standalone monitoring system installed in the build chamber of the EOS M280 machine.System OperationThe monitoring system captures two images per layer of the build process. An example of a pair of these images is shown in Figure 3 below. The first captured image, Figure 3(a), is the deposited powder layer, after re coating, and the second image, Figure 3(b ), is captured after the laser exposure.

The controller then stores the image onto the external hard drive allowing for the required image processing to be completed.Image ProcessingAt present, the images are exported to a desktop PC for ex-situ analysis after the build process is completed. Future work will enable the acquired images to the analyzed in real-time during the AM process. Figure 4 shows a flowchart of the current operations carried out on the images acquired.Export images from machineCrop Images to Area of InterestImageJApply Line Segment and Anomaly DetectionR Statistical SoftwareCompare Sequential ImagesR Statistical SoftwareFigure 4: Flowchart of the image processing operationsSoftware packages ImageJ and R statistical are the used to process the images once they have been exported from the camera controller. ImageJ is an open source image analysis program [29]. A macro script is used to apply a batch crop on all of the images acquired from the camera.

ImageJ is an open source image analysis program [29]. A macro script is used to apply a batch crop on all of the images acquired from the camera. This allows for the regions of interest to be selected and the other areas to be removed. This is shown in Figure 5 below.---- ----- -Acquired image with the cropped build plate area marked in redCropped image of build plateFigure 5: ImageJ cropping tool1394Anomaly detection and line segment detection tools are applied to the cropped images using R Statistical Software [30]. In this paper we focus on the problem of detecting anomalous images of the powder bed pre-laser exposure. Study 1 in the following section discusses an example of the application of these image analysis tools in the PBF process. To aid the identification of process defects, line segment detection is applied, a machine learning technique which identifies locally straight contours known as line segments in digital grey-scale images.

Each pixel of an image has an associated grey-level value and the contours or line segments that are detected by the algorithm are zones of the image where the grey-level is changing suddenly from dark to light or from light to dark [31]. The anomaly detection tool can then be used to detect anomalous counts of line segments in the images. The broader research challenge is to detect anomalous images of the entire build plate as they arrive in real-time and to alert the machine engineer in sufficient time to rectify the issue.Study 1: Powder Short FeedingThe aim of this study is to evaluate the effectiveness of the installed monitoring system and of the newly developed image analysis methodology to identify instances of short feeding of powder, which leads to a lack of powder deposition on the parts being produced. In this study, multiple builds consisting of two sets of four 1 Omm2 cubes were manufactured.

In this study, multiple builds consisting of two sets of four 1 Omm2 cubes were manufactured. The parts were positioned on the build plate in the locations shown in Figure 6. The two sets (Set 1 and Set 2 in Figure 6) of cubes were then cropped as shown in Figure 6, before being analyzed using the line segment detection and anomaly detection scripts.Set 1Set 2Set 1Set 2Acquired Image with cropping regions marked in redFigure 6: Cropping Regions for Set I and Set 2 of partsSet 1 is the control set of samples and are built without any deliberate defects. Set 2 is deliberately subjected to some short-feeding of powder during the build between the build height of 4.5 mm and 6.5 mm. The short feeding will be induced by reducing the volume of powder that is fed to the re-coater blade before a new layer of powder is deposited. This will result in a region on the cubes in Set 2 with a shortage of powder deposited.

This will result in a region on the cubes in Set 2 with a shortage of powder deposited. This lack of deposited powder will result in the re-melting of previous layers causing voids and delamination in the part. The damage to the final part due to the short-feeding is presented in Figure 7. This type of damage can have a detrimental effect on a part, if not detected within two or three layers of the start of the short-feeding.1395Figure 7: Set 2 parts - with damage due to short-feeding highlightedThe short-feeding was applied by reducing the amount of powder that is deposited on the build plate. This is controlled by the re-coating percentage, which can be explained as a percentage of the volume of the build area. Typical recoating parameters are 120%; meaning that 120% of the volume in the build area (i.e.120% layer thickness by the area of build platform) is re-coated. This ensures that the entire build plate is covered.

This ensures that the entire build plate is covered. The short feeding of powder was created through the reduction of the re-coating parameters to 80%, for layers between the heights of 5 mm and 6 mm. This created an area on Set 2 of the cubes which received no powder during the recoating stage.For example, images that capture unusual lines or patches of powder can be seen on the build plate. Figure 8(a) shows a typical image where a consistent layer of powder was deposited correctly on Set 2 parts prior to the introduction of short-feeding. Figure 8(b) shows and example of where short-feeding has occurred on these parts. This can be seen as an area on the part where there is no deposited powder, which exposes the previously melted layer.

This can be seen as an area on the part where there is no deposited powder, which exposes the previously melted layer. Typically, two to three problem layers can cause a defect in the built part.(a) New Powder Layer - no short feeding(b) New Powder Layer- short feedingFigure 8: In-situ camera images showing (a) an example of correct powder distribution on the cubes in Set 2 after recoating and (b) an example of short-feeding resulting in insufficient powder being deposited on the cubes in Set 2.1396The line segment detection tool has been used to identify this issue of short feeding, shown in Figure 8(b ). When the short feeding of powder begins, the area with a lack of powder will show an increase in the number of line segments detected. This is indicated by a spike in the plot of the number of line segments per image, as shown in Figure 9.

This is indicated by a spike in the plot of the number of line segments per image, as shown in Figure 9. The anomaly detection tool is then used to track any deviation in the quantity ofline segments detected and then inform the machine operator of the increase in the line segments detected.VI +-' C Q) E bO Q) VI Q) C..,_""C Q) +-' u Q) +-' Q) ""C20 --0I... Q) ..0 E :l zo-J~lui.u~.liliu1w.1500175020002250Image IDFigure 9: Line segment detection tool output with cooresponding Set 2 build images.In Figure 9 it is clear from the number ofline segments detected where a defect has begun to develop. The spike in the graph occurs 5.4 mm (Image ID 2041) into the build process. The deliberate short feeding of powder has reduced the quantity of powder deposited on the build plate, leaving an area on the parts without new powder. This exposes the previously melted layer which is then re-melted by the laser.

This exposes the previously melted layer which is then re-melted by the laser. This issue of short feeding can be rectified by operator intervention, if noticed in time. The operator can, once the defect has been identified, increase the quantity of powder fed to the re-coater blade. This intervention was applied in the study after 6.0 mm (Image ID 2101) of the build by increasing the amount of powder fed to the re-coater blade. This is shown in Figure 9 by a decrease in the line segments detected as the powder deposition recovers (Image ID 2130).ConclusionsIn this paper, a standalone monitoring system for the PBF process is presented. An accompanying image processing methodology has been developed, which can identify defects in the powder deposition stage of the DMLS process. The algorithm is computationally efficient and works well for the application presented. The images collected for Study 1 were analyzed post-build using R statistical software on a remote server.

The algorithm is computationally efficient and works well for the application presented. The images collected for Study 1 were analyzed post-build using R statistical software on a remote server. However, the method is efficient enough to use in real-time and work is ongoing to enable this process on site. With this facility an anomaly detection algorithm will be used to detect anomalous counts of line segments in the images.Future work will expand the capability of the algorithm to enable defect detection in the images post-laser exposure, as well as enabling the defect detection to operate in real time with the process alerting the machine1397operator of any identified defects.

Thus, allowing the operator to alter the process as required to ensure the successful outcome of the AM build, reducing scrap costs for the process.Acknowledgements This publication has emanated from research supported in part by a research grant from Science Foundation Ireland (SFI) under Grant Number 16/RC/3872 and is co-funded under the European Regional Development Fund. The authors would like to thank Irish Manufacturing Research for their funding and support of this research.ReferencesA. T. Sutton, C. S. Kriewall, M. C. Leu, and J. W. Newkirk, "Powder characterisation techniques and effects of powder characteristics on part properties in powder-bed fusion processes," Virtual Phys. Prototyp. , 2016, vol. 12, no. 1, pp. 3-29. J. H. Tan, W. L. E. Wong, and K. W. Dalgarno, "An overview of powder granulometry on feedstock and part performance in the selective laser melting process," Addit. Manuf , 2017, vol. 18, pp. 228-255. [3] M. Mani, B. M. Lane, S. C. Feng, S. P. Moylan, A.

This paper describes a reference framework for AM data integration, named AMIF, and the design of the testbench based on AM Integration Framework (AMIF) for testing the integration of in-process data acquisition, real-time feature extraction, process control, and predictive models under a data management system. A specification of this testbench is developed to manage and stream voluminous data captured by high-speed cameras and performing data analytics using common information models and functional interfaces. The integration of the data, models, and computer tools sends operational decisions to an AM machine in real time. On top of the real-time control functions, AM data integration with MES and ERP systems is also included using a high-performance data warehouse for long-term data archiving and metadata management. The architecture of this testbench is illustrated in this work.

The architecture of this testbench is illustrated in this work. AMIF can guide AM practitioners and system integrators to build their integrated AM manufacturing systems for production. The NIST AM testbench’s plug-and-play features allow both internal and external researchers and developers to assess the effectiveness of their individual data models, data analytics, and decision-making algorithms on the systems engineering level.1. IntroductionAdditive Manufacturing (AM) demonstrates its capability of building complex geometry and customized products. AM is a strong contender to the current manufacturing methods with many successful applications, including fabrication of biomedical implants and aerospace engine parts [1]. Laser Powder Bed Fusion (LPBF) is a kind of metallic AM technology that employs lasers to fuse layers of metallic powder at specific energy levels and speeds.

Laser Powder Bed Fusion (LPBF) is a kind of metallic AM technology that employs lasers to fuse layers of metallic powder at specific energy levels and speeds. This technique builds near-net-shape parts using a layer-by-layer approach.LPBF can significantly reduce processing steps to build complex parts; however, a long list of variables may affect the quality of AM parts, such as process parameters, environmental conditions, and material status. For example, the microstructure and mechanical features of the AM build parts are deeply dependent upon the process parameters, such as laser power, and scan speed. By fine-tuning the process parameters, the mechanical characteristics of the building samples can be enhanced [2]; however, when multiple build samples are created using the same2166process parameter conditions, the quality of each finished sample is dissimilar with each other.

This result can be due to different factors, such as coating homogeneity, airflow stability, and fusion capability [1]. Also, building a delicate AM part usually takes a long time to complete the process. Evaluating the quality of the building process in real-time can assist the in-process decision-making and improve productivity.To better evaluate an LPBF process, the in-situ measurements can provide rich information in different length and time scales. Different types of sensors are added to AM machines and these sensors generate large volumes of structured or unstructured measurement data at high frequencies. Successful integration of data and their analysis will enable effective in-process monitoring and real-time control [3].At the same time, the maturation of AM into broad production and industrialization requires an expanded notion of integration for both AM systems and AM data.

System and data integration includes other types of traditional manufacturing systems, manufacturing operation functions, and broader business processes across AM value chains, which are defined by ISA 95 Level 3 and 4 functions. Figure 1 shows the function hierarchy of manufacturing systems defined in ISA 95.Figure 1. ISA 95 Standard [4]ISA 95 is an international standard from the International Society of Automation for developing integration interfaces between enterprise and control systems. Level 0 defines the physical production processes. Level 1 defines the activities included in sensing and manipulating the physical processes. Level 2 defines the activities of monitoring, supervising, and controlling the physical processes. Level 3 defines the activities of the workflow, such as the process of manufacturing products from raw materials to finished products.

Level 3 defines the activities of the workflow, such as the process of manufacturing products from raw materials to finished products. Level 4 defines the business- related activities required for Level 3.For AM process manufacturing systems, integration is needed to automate engineering workflow and improve decision-making across all the function levels of ISA 95 covering a complex set of engineering domains, such as design, material, process, machine, and manufacturing. However, the ISA 95 function architecture is not a sufficient guide to model the system and data integration for AM industrialization. The desired system requires high-speed data processing and integration for advanced data analytics, including the use of machine learning or deep learning for process planning, control, and part qualification. Figure 2 captures the system integration scope for AM and highlights its specific need for data integration (in blue fonts).2167Figure 2.

Figure 2 captures the system integration scope for AM and highlights its specific need for data integration (in blue fonts).2167Figure 2. AM Specific Integration RequirementsIn this research, we developed an AM integration framework that leverages ISA 95 addressing the specific high-speed data integration and analytics needs. Based on this reference architecture, we designed an AM integration testbench that can be used to test various communication protocols, information models, and advanced AI functions for AM development production and deployment. This testbench covers data collection, management, and analysis to automate offline engineering functions and data flow in real-time or near real-time. The testbench is designed for both internal and external use.This paper is organized in three sections. In section 2, we describe the AM Integration Framework (AMIF) as a reference architecture for any modern manufacturing systems including AM processes.

In section 2, we describe the AM Integration Framework (AMIF) as a reference architecture for any modern manufacturing systems including AM processes. In section 3, we present an implementation architecture for the NIST AM integration testbench, which employs an AM emulator instead of a real AM machine. We also listed some potential communication protocols, information models, and advanced analytical functions. In section 4, we explain the use of AM big data storage and display the test results for determining the current data streaming capabilities.2. AMIF Leverage on ISA 95We leverage ISA 95 to define an AMIF for specifying the common AM data exchange scenario between various functions or applications to offer a generic reference framework for AM data integration.The AMIF we propose is to leverage the ISA 95 function architecture but only define three different levels for integration purposes.

In Figure 3, the AMIF displays the process relationship of how data, information, and events work between each function. Level 1 defines the main functions provided to AM machines, which corresponds to ISA 95 Level 0 and 1. Between the AMIF Levels 1 and 2, we have a stage called “real-time” monitoring & control. This stage is necessary to support third-party advanced monitoring add-ons and functions faster than edge computing can offer; for example, FPGA is required to support this kind of high processing speed. Level 2 defines the activities including near real-time monitoring and control, which corresponds to the ISA 95 Level 2. This level relies mostly on edge computing. Level 3 defines activities,2168including data management, manufacturing operation, and enterprise application, which corresponds to ISA 95 Levels 3 and 4.2.1 AMIF DiagramFigure 3.

Level 3 defines activities,2168including data management, manufacturing operation, and enterprise application, which corresponds to ISA 95 Levels 3 and 4.2.1 AMIF DiagramFigure 3. AMIF Diagram2.1.1 Functions in AMIF Level 1 - AM Machine/EmulatorThe AMIF Level 1 captures common functions the AM machines are designed with. Process Monitoring is for monitoring the process with high and low speed to collect the raw data for process analysis. Process Control is for controlling machine operation or monitoring devices with control systems based on the original process design command or command generation function during feedback control. Machine Information is for outputting or updating the information of the current machine status (Event-based).

Machine Information is for outputting or updating the information of the current machine status (Event-based). Each edge-calculation function should activate according to the current process status, so an operating system dedicated to function communication must be installed for near real-time feedback control.2.1.2 Real-Time Monitoring & Control StageHigh-Speed Data Acquisition is needed for capturing and processing data from advanced in-situ monitoring systems in AM machines, facilitating both high-speed operation and real-time control. Feature Extraction is for analyzing the data from high-speed data acquisition and discovering useful features for anomaly detection. Anomaly Detection is based on the feature extraction results. This is used to detect abnormalities in each sampling data. Command Generation is based on the anomaly detection result. This is used to generate the command for2169high-speed feedback control.

This is used to detect abnormalities in each sampling data. Command Generation is based on the anomaly detection result. This is used to generate the command for2169high-speed feedback control. Feedback Control is for transmitting the command to the process control function located in the AM machine level for high-speed process compensation.2.1.3 Functions in AMIF Level 2 – Near Real-Time Monitoring & ControlData Aggregation is for carrying out corresponding data grouping according to different analysis requirements of scale. Aggregated Feature Extraction is based on the aggregated data performing feature calculation allowing the acquisition of useful features at a certain scale/resolution. Data Streaming is for uploading the aggregated data to an online data archive service located in the data management level. Quality Estimation is for using an AI/ML model for online quality estimation. The model is created offline in a model creation function located in the AMIF Level 3.

Quality Estimation is for using an AI/ML model for online quality estimation. The model is created offline in a model creation function located in the AMIF Level 3. Updated Command Generation is based on quality estimation results. This level generates the command for event-based feedback control. Machine Operation Control is for getting the AM machine information and controlling the process according to the different feedback control requirements. The function to update the machine information, located in the AM machine/emulator level, is based on the event-based command of feedback control.2.1.4 Functions in AMIF Level 3 - Data Management/Manufacturing Operation/Enterprise ApplicationBig Data Repository is for archiving AM raw data. Metadata Store is for saving and managing the metadata of big data objects or building pedigree information. Data Visualization is for visualizing the processed data of functions in AMIF Level 2.

Metadata Store is for saving and managing the metadata of big data objects or building pedigree information. Data Visualization is for visualizing the processed data of functions in AMIF Level 2. Model Creation Offline is for creating the AI/ML model for online quality estimation, including feature selection, model training, and validation for different prediction targets or purposes. Product Lifecycle Management (PLM) is an information management system that integrates data, processes, and business systems. By estimating AM part quality, we can optimize the AM product geometry design. Machine Monitoring is for monitoring the machine information, including process operation, parameters, and sensors. Quality Management System (QMS) is a collection of business processes focused on consistently meeting customer requirements and enhancing their satisfaction. Build Planning uses data analysis results to optimize the building strategy and design.

Build Planning uses data analysis results to optimize the building strategy and design. After build optimization, it generates the pre-program command.3. Testbench Design for AM Operation and ControlThe functional components in Figure 4 are essential to facilitate comprehensive testing of AM data integration scenarios. The NIST AM Data Integration Testbench implements the AMIF function architecture at three levels. However, instead of working directly on AM machines, which may result in the disruption of AM processes, the AM testbench at NIST utilizes an AM emulator. The AM Emulator level focuses on AM process event generation and sends the in-situ monitoring images out with a high-speed camera protocol. The Edge Computing level focuses on high-speed image acquisition and other necessary functions, such as streaming or decision-making. This level also implements a machine operation control. The Cloud level is for saving and managing raw data and metadata.

This level also implements a machine operation control. The Cloud level is for saving and managing raw data and metadata. These data can be used by manufacturing operations and management functions2170or enterprise functions for decision-making. In this section, we will introduce the main test modules for different levels of data integration for AM operation and control.Figure 4. Testbench Design to Optimize AM Data Integration Performance3.1 AM Process Event GeneratorThe event generator is for simulating the occurrence of events and utilizing a publish/subscribe interface to communicate with the functions of AM operation control. A working AM operation control follows the AM process step allowing the function in edge computing to work at the right time.

A working AM operation control follows the AM process step allowing the function in edge computing to work at the right time. If the decision-making needs to do a certain AM process for feedback control, the system utilizes this event generator to generate that event.For example, when getting the ending event of the powder coating in Figure 5, this mechanism will inform the global camera to take the global image. After the function analyzes the image, a decision is sent back to the event generator to decide if we need to perform a recoat to control the coating quality. After the laser melting process, we estimate if the surface roughness is under the desired specification and send the decision to the event generator for the next step [1][5].Figure 5. AM with Feedback Control21713.2 AI-Based Image GeneratorFigure 6 illustrates the input and output of the AI-based image generator.

AM with Feedback Control21713.2 AI-Based Image GeneratorFigure 6 illustrates the input and output of the AI-based image generator. The generator takes one-dimensional process parameters, such as laser power, scan speed, and direction along with random noise as its inputs, to produce the corresponding melt-pool image (MPI). This generator is built upon deep learning generative models, including generative adversarial networks and the Diffusion Model. Within the workflow of the AM emulator, the image generator emulates the MPI based on prescriptions of the AM process event generator.Figure 6. AI-Based Image Generation3.3 Camera Simulator with High-Speed Communication ProtocolThe purpose of using a high-speed camera in AM is to study material behavior, process monitoring, and quality control.

By recording the process in real-time at high frame rates, manufacturers can closely observe potential defects and perform in-process feedback control.The camera simulator can generate a high-speed video stream or test patterns for testing frame grabbers or vision/imaging systems. It allows developers to quickly prototype their vision and imaging systems without the need for physical cameras. It also enables the testbench to evaluate the performance of the high-speed camera interface, such as CoaXPress and CameraLink, in transmitting image data. CoaXPress reaches 50 Gbps by using four lanes which enables a higher data rate than CameraLink. Table 1 shows the comparison between CoaXPress and CamerLink [3].If the camera requirement generates less than 125 MB/s (1 Gbps), both Gig-E and USB3 Vision can be used. For example, each grayscale image is 120 by 120 pixels, and the bit depth is 8 bits.

For example, each grayscale image is 120 by 120 pixels, and the bit depth is 8 bits. To capture 10,000 images per second, the data rate needs to achieve 137 MB/s (1.07Gbps) [6][7].Table 1. Comparison between CoaXPress and CamerLink [6][7]InterfaceCoaXPressCamera linkConnectorBNCMDR or SDRCable Length30 to 212 meters1 to 10 meters2172Data RateReal Time Trigger12.5 Gbps per lane 2.0). (CoaXPress Uses lanes four which can reach 50 Gbps. Yes255 Mbps to 850 Mbps, depending on the specific implementation mode. (Base, Medium, Full, Extended-Full) Yes3.4 Event-Based Communication ProtocolMachine monitoring is crucial for a variety of applications. Preventive maintenance, optimizing productivity, quality estimation, or feedback control all rely on process and sensor data transmitted by machines. OPC UA and MQTT are common communication protocols used in industrial settings for data exchange and interoperability. Table 2 shows the comparison between them.

OPC UA and MQTT are common communication protocols used in industrial settings for data exchange and interoperability. Table 2 shows the comparison between them. If the message includes structure data, the OPC UA is more suitable. If the message is similar to a trigger signal, MQTT is the better choice for the lightweight case.Table 2. Comparison between OPC UA and MQTTProtocolOPC UAMQTTUse caseProvide security and reliability. Widely used in complex industrial automation, PLC communication, etc.Lightweight messaging protocol. For resource- constrained environments, like IoT devices, telemetry applications, or various low- bandwidth conditions.Communication ModelsClient-Server Architecture Publish-Subscribe mechanism.

For resource- constrained environments, like IoT devices, telemetry applications, or various low- bandwidth conditions.Communication ModelsClient-Server Architecture Publish-Subscribe mechanism. (Introduced in OPC UA 1.04).Publish-Subscribe mechanism.Overhead Read/Write OperationsHigher overhead Read and writeLower overhead Read and write3.5 AM Data Streaming and ManagementAutomating the process of uploading data to the cloud in real time during AM process is important. We can utilize the cloud services to set up cloud storage for data integration. NIST has already developed a collaborative AM data management system, providing data storage in clouds structured by an AM lifecycle data schema [8].

NIST has already developed a collaborative AM data management system, providing data storage in clouds structured by an AM lifecycle data schema [8]. In this data integration testbench, we will also build a similar AM data management system to simulate the data streaming and archiving application as well as explore the different methods of storage and evaluate their capabilities.2173The process of AM produces large amounts of data for long-term archiving and future data analysis. AM big data may include CAD model files and in-process melt-pool images. The new framework desires two different types of storage systems for archiving AM big data and metadata separately. Currently, the Additive Manufacturing Materials Database (AMMD) [8], a NoSQL MongoDB document data warehouse developed by NIST, is the data storage used to store structured and unstructured AM metadata.

However, there are limitations with the current storage system that may affect the potential speed and performance of the AM testbench.3.5.1 AM Metadata StoreAs the quantity of generated AM data grows, searching for a specific dataset in a big data repository becomes increasingly difficult. This system requires cooperation with a database capable of holding structured metadata that provides details for data objects stored in the cloud. An acceptable option for a metadata store would be a document database, as it allows highly structured documented data architecture for immutable AM metadata. Another option worth exploring is a graph database, which focuses on graphing node entities and edge relationships.3.5.1.1 Document StorageThe current AMMD [8] storage system stores AM metadata in XML format and follows an XML Schema Definition (XSD) data structure which partitions XML documents into 9 types.

Each document type has its own validated structure layout and are linked together by identification values. These databases can handle large quantities of data storage and are flexible to small changes to the schema structure. What makes the system limited is the difficulty in establishing complex relationships between metadata located in different XML documents. Each separate document establishes a connection by the document’s identifier; however, the relationships between subsets of metadata within each document are not clearly specified. This makes data query a slightly more tedious task.3.5.1.2 Graph StorageGraph databases serve to be an efficient AM data analysis tool with easily navigable node and edge relationships. In comparison to document-based databasing, the graph makes use of relationships between node entities as the primary way to draw connections between data in the entire graph view.

In comparison to document-based databasing, the graph makes use of relationships between node entities as the primary way to draw connections between data in the entire graph view. This method is great for data analytics because it consolidates data, enabling quicker traversal between the node entities and edge relationships; however, the database’s performance is reliant on a well-defined data ontology. An unnecessary or undefined relationship in the ontology may result in a separation of important information between two related nodes.Much remains unknown about integrating AM data with graph databases. We are currently investigating the metadata storage and query capabilities of Deep-Lynx [9], an open-source graph data warehouse developed by Idaho National Laboratory (INL), to obtain a closer understanding of the compatibility of node and edge architecture along with the AMIF.21744.

AM Big Data Streaming and ArchivingTo determine if a storage system fits our criteria, we emulate the AM data generation and streaming processes to test a database’s features and capabilities. The original system utilizes batch uploads to archive all types of AM data, but this method is an inefficient use of time. By streaming AM data during the AM process, we can improve the speed of archiving data, making data readily available for analysis at any time.4.1 Utilize S3 Cloud StorageStreaming and archiving AM big data requires a high-volume file storage system compatible with fast and continuous upload. Amazon Simple Storage Service (S3) is the selected cloud object data lake storage which offers scalable file storage, configurable security access, and several API software development kits. The concern of using this service is the lack of understanding about the S3 capabilities and internet data transfer rate.

The concern of using this service is the lack of understanding about the S3 capabilities and internet data transfer rate. A vital component of near real-time data streaming is a consistent high-speed connection between local AM machines and cloud services. Slower connections result in a growing backlog of data queued for upload. To ensure timely processes, an AM in-process simulation is constructed to emulate the procedure of generating image data, compressing each package, and streaming the data to the designated S3 bucket destination.4.2 Upload Speed Test DatasetThe AM in-process simulation consists of artificial parameters representing the desired and testing upload speed configurations. The build is a 12-part model each constructed with 250 layers. One package contains sample image data for each part of each layer resulting in a total of 3000 packages produced per test run.

The build is a 12-part model each constructed with 250 layers. One package contains sample image data for each part of each layer resulting in a total of 3000 packages produced per test run. Regarding the internet connection, the simulation uses an RJ45 ethernet cable to connect local AM processors to the AWS cloud services. The simulation compresses each package into a ZIP file since compressed data results in a quicker internet transfer rate. In Table 3, if we divide the data size per package by the compressed data size per package, the compressed package is estimated to be ¼ of the full package size.Table 3. Upload Speed Test Dataset Base ParametersConfiguration # of Layers # of Parts Total # of Packages # of Images per Package Single Image Size Data Size per Package Compressed Data Size per Package (ZIP)Data 250 12 3000 1500 43.3 KB 64.881 MB 16.92 MB2175The setting for the camera frame rate is around 4000 FPS.

During the time frame of layerscanning and recoating, image data is captured and streamed to the cloud.Table 4. Upload Speed Test Scan and Recoat Time ParametersData 4.8 seconds 5 seconds Total Time for Minimum Buffer 9.8 secondsConfiguration Scan Time per Layer Recoat Time per Layer4.3 Desired Speed CalculationEstimating the minimum upload speed requirement involves creating mock configuration parameters that imitate an actual AM process. Results of the upload speed simulation that fall short of the minimum requirement could end up with an undesirable stockpile of data files queued for cloud transfer.Table 5. Minimum Desired Upload Speed Base ParametersConfiguration # of Layers # of Parts Data Size per Package Scan Time per Layer Recoat Time per LayerData 250 12 60 MB 4.8 seconds 5 secondsIn Table 6, the calculated desired upload speed for one package is obtained by dividing data size per layer by total time per layer.

The actual and compressed data size from the upload speed testing dataset is rounded to a ¼ compression ratio, so the estimated desired upload speed for the compressed package is acquired by dividing the desired upload speed by 4.Table 6. Minimum Desired Upload Speed CalculatedConfiguration Total Time per Layer Data Size per Layer Desired Upload Speed Desired Compressed Upload Speed (ZIP)Data 9.8 seconds 720 MB 73.469 MB/s 18.367 MB/s4.4 Upload Speed Test Results2176The test results will be compared to the estimated desired upload speed. If the upload speed test results exceed the desired speed, we can say the AM data transfer to the S3 bucket will be able to keep pace with AM machine data generation. In Table 7, the average data upload speed displays the theoretical transfer speed of the data size within a compressed package, not the upload speed of each individual data file.Table 7.

In Table 7, the average data upload speed displays the theoretical transfer speed of the data size within a compressed package, not the upload speed of each individual data file.Table 7. Local to S3 Upload Speed Test ResultsConfiguration Avg ZIP Upload Time per Package Avg ZIP Upload Speed Avg Data Upload SpeedData 0.446 seconds 37.937 MB/s 145.459 MB/sAfter multiple simulations, the accumulated data averaged a ZIP package transfer rate of 37.937 MB/s meaning the data inside each package uploaded at a speed of around 145.459 MB/s. The test results more than doubled the minimum required upload speed.Table 8. Comparison between Desired and Test Result Upload SpeedUpload Speed Compressed Package (ZIP) Data inside PackageDesired 18.367 MB/s 73.469 MB/sTest 37.937 MB/s 145.459 MB/sIt is important to note that the upload speed test simulation does not contain each layer’s time configurations for laser printing, powder filling, build plate descent, and other factors.

Theoretically, adding these time parameters would significantly lower the estimated desired upload speed making the current simulation test results even more qualified for near real-time AM data streaming.4.5 Combining Image and Metadata Streaming ProcessesThe AM data streaming simulation runs simultaneously with each other. When generated, data is recognized and uploaded to the specified endpoint. What makes this process problematic is how both streaming methods are run asynchronously from each other. This means data in Amazon S3 is not linked with metadata within the graph database. A potential solution could be to send an identifier metadata value within the S3 upload response back to the local computer. Then, a separate process will send an API request containing the S3 object identifier to the graph database and update the corresponding node metadata in the graph view.21775.

Then, a separate process will send an API request containing the S3 object identifier to the graph database and update the corresponding node metadata in the graph view.21775. SummaryAMIF is designed as a framework for AM system developers to integrate process data and sensor signals. The NIST AM data integration testbench is a use case that emulates a real AM system process for testing the interoperability of computer tools and evaluating the machine operation control. The testbench also demonstrates its capability to handle high-speed, high- volume in-process data in an autonomous workflow. This testbench can serve as a standard platform to test the integration of data flow and the pipeline of software operations. The goal is to integrate industrial cases on the testbench to evaluate the various communication protocols, information models, and numerical simulations.

The goal is to integrate industrial cases on the testbench to evaluate the various communication protocols, information models, and numerical simulations. Also, utilizing this testbench can enable testing advanced LPBF process monitoring and predictive algorithms for near real-time or real-time data analysis and process control.Disclaimer and AcknowledgementCertain commercial equipment, instruments, or materials identified in this paper are not intended to imply recommendation or endorsement by the National Institute of Standards and Technology, nor is it intended to imply that the materials or equipment identified are necessarily the best available for the purpose.The authors would like to thank Dr. Ho Yeung for the support and information sharing for the testbench development. We also feel thankful for the support and the insightful discussions provided by Dr. Paul Witherell and Dr. Brandon Lane at NIST.Reference[1] M. Adnan, H. -C. Yang, T. -H. Kuo, F. -T. Cheng and H.

Paul Witherell and Dr. Brandon Lane at NIST.Reference[1] M. Adnan, H. -C. Yang, T. -H. Kuo, F. -T. Cheng and H. -C. Tran, "MPI-Based System 2 for Determining LPBF Process Control Thresholds and Parameters," in IEEE Robotics and Automation Letters, vol. 6, no. 4, pp. 6553-6560, Oct. 2021.[2] G.Tapia, A. H. Elwany, and H. Sang, “Prediction of porosity in metal-based additive manufacturing using spatial Gaussian process models,” Additive Manuf., vol. 12, pp. 282–290, 2016.s[3] Perisic, Milica, Yan Lu, and Albert Jones. "In-Process Data Integration for Laser Powder Bed Fusion Additive Manufacturing." International Design Engineering Technical Conferences and Computers and Information in Engineering Conference. Vol. 86212. American Society of Mechanical Engineers, 2022.[4] ISA95, Enterprise-Control System Integration, https://www.isa.org/isa95/, Accessed on May 20, 2020.[5] Yang, Haw-Ching, et al.

Vol. 86212. American Society of Mechanical Engineers, 2022.[4] ISA95, Enterprise-Control System Integration, https://www.isa.org/isa95/, Accessed on May 20, 2020.[5] Yang, Haw-Ching, et al. "An intelligent metrology architecture with AVM for metal additive manufacturing." IEEE Robotics and Automation Letters 4.3 (2019): 2886-2893.[6] CoaXPress Standard, https://www.coaxpress.com/; [accessed on August 29, 2023].2178[7] CameraLink HS Standard, https://www.automate.org/a3-content/vision-standards-camera- link-hs; https://knowledge.ni.com/KnowledgeArticleDetails?id=kA00Z000000P7rmSAC; [accessed on August 29, 2023].[8] Lu, Yan, Paul Witherell, and Alkan Donmez. "A collaborative data management system for additive manufacturing." International design engineering technical conferences and computers and information in engineering conference. Vol. 58110.

"A collaborative data management system for additive manufacturing." International design engineering technical conferences and computers and information in engineering conference. Vol. 58110. American Society of Mechanical Engineers, 2017.[9] Deep-Lynx, https://github.com/idaholab/Deep-Lynx; [accessed on August 31, 2023].2179.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDevelopment of a Thermoplastic Biocomposite for 3D PrintingJohn Obielodan,* Joshua Helman, and Andrew GrumblesDepartment of Mechanical Engineering, University of Wisconsin-Platteville *Corresponding author: obielodanj@uwplatt.eduAbstractOrganosolv lignin, a natural cross-linked phenolic wood polymer is a by-product of thepulping process in the paper industry. It is a renewable organic natural product with potential application in many areas. It has attractive properties that make it a potential candidate for fabricating useful parts using 3D printing. Also, polylactic acid (PLA), a biodegradable thermoplastic derived from renewable sources is widely used in 3D printing polymer parts.

Also, polylactic acid (PLA), a biodegradable thermoplastic derived from renewable sources is widely used in 3D printing polymer parts. This work seeks to study the technical viability of extruding different blends of PLA and organosolv lignin into filaments for 3D printing useful objects. Filament extrusions using different blends were evaluated. Also, the mechanical properties of printed test samples are presented. Viable blends of the biocomposite for 3D printing has the potential to provide an added-value to lignin for expanded use in many applications.Introduction3D printing has revolutionized how products are made and has become a benchmark technology that enables innovative materials development for a wide range of applications. Polymer-based 3D printing has mostly depended on petroleum-based materials like Acrylonitrile Butadiene Styrene (ABS), Polycarbonate (PC), Nylon and many others with significant long- term health and environmental consequences.

To address these concerns, biomaterials having comparable qualities and properties are being sought as alternatives in many applications. Polylactic Acid (PLA) [1-3], a biodegradable and environmentally benign thermoplastic derived from renewable sources has been available as alternative 3D printing material for many applications. Some biocomposites made of a blend of laywood, laybrick, bamboo, straw, and carbon fiber with PLA and some other biopolymer are also commercially available for divers applications [1, 4].Lignin is one of the most abundantly available organic polymers on earth, second only tocellulose. It is a product of the pulping process in the paper industry and biofuel production [5- 6]. Although, it is used in some applications including the manufacture of vanillin, animal feed, dye dispersants, resins and cleaning chemicals, it has mostly been used as fuel for the recovery of its energy content in the paper industry.

The annual production of lignin is estimated to be 50 million tons, with only 2% being used for applications other than combustion and energy production [6]. There are different types of lignin depending on processing such as kraft lignin and organosolv lignin. They are cheap materials for which high-end applications are being sought. Some earlier work have on attempted to extrude a blend of kraft lignin with PLA for 3D847(a)printing. Gkartzou, et al. [7], presented their work on the blends of kraft lignin with PLA with limited success. The blend was extruded successfully into filament at 5% kraft lignin for 3D printing experimental samples. The work highlighted challenges in achieving homogenous blends due to agglomeration of the kraft lignin in the PLA matrix that could result in nozzle clogging during printing. The clumps constitutes a limiting factor to attaining high weight fraction of lignin in the PLA matrix.

The clumps constitutes a limiting factor to attaining high weight fraction of lignin in the PLA matrix. The blends were found to be more brittle than the PLA matrix such that a 20 wt% composition experimental samples made by other methods could not be cut into the standard tensile specimens successfully. The present work explores the potential of organosolv lignin as filler material in PLA matrix for extrusion-based 3D printing applications.Experimental ProcedurePulverized PLA pellets used as the matrix material for the biocomposite polymer in thiswork were obtained from Filabot. They were blended in various proportions with organosolv lignin and extruded into filaments. Scanning electron micrographs of the PLA pellets and organosolv lignin particles are shown in Fig. 1. Before blending, the PLA pellets and lignin powders were dried in a Barnstead Lab-line vacuum oven for at least 48 hours at 45oC to remove moisture contents.

1. Before blending, the PLA pellets and lignin powders were dried in a Barnstead Lab-line vacuum oven for at least 48 hours at 45oC to remove moisture contents. The blends were made to obtain visually uniform dispersion of the lignin in PLA. Filabot EX2, a Filabot product was used for the filament extrusions. A spooler, obtained from the same vendor was used to wind the filaments on spools and for controlling filament tension for dimensional control. A Filabot airpath wind blowing device was placed in-between the extruder and spooler to provide cooling air flow for heat dissipation from the filament. A digital measuring machine was attached to the spooler for real-time diameter measurement to enable dimensional control.

A digital measuring machine was attached to the spooler for real-time diameter measurement to enable dimensional control. The extrusion temperature ranges from 170oC to 175oC.Figure 1: SEM of (a) PLA and (b) organosolv ligninExtruded filaments were sealed in plastic bags with silica gels placed inside to preventmoisture absorption before and after use for 3D printing of test samples. Creality 10-S 3D printer was used to 3D print tensile test samples with five replicates for plain PLA, 1 wt%, 5 wt%, 10 wt%, 15 wt%, 20 wt%, and 25 wt% lignin contents in this work. Cura slicing software was used848for all fabrications with zig-zag raster tool path and settings aimed to achieve 100% 3D print density. Nozzle temperatures varied from 190oC to 210oC depending on composition, while the bed temperature was maintained between 50oC and 60oC. Test samples were 3D printed and tested according to ASTM D638 standard.

Test samples were 3D printed and tested according to ASTM D638 standard. Microstructure of some samples were observed using scanning electron microscopy (SEM).Results and DiscussionThe filament extruded for all compositions are comparable to those that are commerciallyavailable in smoothness. Blended materials exhibited very good extrudability and flowability with no observable agglomeration of the lignin content. Dimensional consistency was achieved within 1.77±0.06 mm. The results of tensile specimens show a decrease of tensile strengths with increasing amounts of lignin content. Figure 2 shows 3D printed PLA has a tensile strength of 57.4 MPa, while those with 10%, 15%, 20%, and 25% lignin has strength values of 45.3 MPa, 40.6 MPa, 34.4 MPa and 32.9 MPa. The strengths wanes with increasing lignin content because there is no effective load transfer between the constituent materials.

The strengths wanes with increasing lignin content because there is no effective load transfer between the constituent materials. These strength values are however, within the range or superior to disclosed values for some 3D printed engineering grade polymers produced by Stratasys [8-9].There’s no statistical difference between stiffnesses of different sample compositions as can be seen in Fig. 3. However, Fig. 4 show a pattern of improved ductility with increase lignin content. Lignin concentrations from 15 wt% to 25 wt% yielded significantly better ductility than plain PLA.

5 reveals the blend qualities of the constituent powdered materials, 3D printing qualities, and pattern of failure. Fractured surfaces show the constituent materials were well blended prior to and during the extrusion processes. There is no visual evidence of agglomeration that could have affected the extrusion qualities. Although the fabrication parameters required to obtain 100% density were selected, some incidences of porosities characteristic of extrusion-based 3D printing can be observed.

Although the fabrication parameters required to obtain 100% density were selected, some incidences of porosities characteristic of extrusion-based 3D printing can be observed. The effects of these porosities on the tensile strengths of the tested samples can only be verified when higher density samples can be made using better software and/or hardware, or if a post processing step that closes the porosities is applied.850(a)(b)(c)Figure 5: Representative SEM fractographs of tensile samples for (a) 10 wt% lignin, (b) 15 wt% lignin, and (c) 20 wt% lignin contents.ConclusionsThe range of the PLA/Lignin biocomposite material blends used for fabrication in this work demonstrate the viability of lignin for 3D printing applications. The results surpassed the limitations presented in earlier work using kraft lignin [7]. The blends of up to 25 wt% demonstrated good flowability during extrusion and 3D printing.

The results surpassed the limitations presented in earlier work using kraft lignin [7]. The blends of up to 25 wt% demonstrated good flowability during extrusion and 3D printing. The tensile strengths obtained are comparable to petroleum-based material properties under ambient temperature applications. The stiffness is virtually unaffected by increase in lignin content. There is improvement on the ductility of the biocomposite blends over plain PLA. Results for higher blend compositions of the biocomposite beyond the scope of this work will be presented elsewhere. The use of organosolv lignin for 3D printing has the potential for repurposing it for high-value niche applications that could result in lowering the cost of materials for extrusion-based 3D printing.851AcknowledgmentThis research was supported by WiSys and UW System applied research funding programs. The views expressed herein are those of the authors and are not necessarily those of WiSys or UW System.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDevelopment of an acoustic process monitoring system for selective laser melting (SLM)N. Eschner*, L. Weiser*, B. Häfner*, G. Lanza*wbk – Institut für Produktionstechnik, Karlsruher Institut für Technologie, 76131 Karlsruhe1 AbstractThe current selective laser melting (SLM) process lacks both process quality and reproducibility. Recent research focuses on the integration of optical measuring technology, but acoustic sensors also seem promising. Initial results on acoustic methods show their suitability. The further processing of the data still shows difficulties, mostly due to the high sample rate. In this work a concept for an acoustic process monitoring system is developed and integrated into the process.

In this work a concept for an acoustic process monitoring system is developed and integrated into the process. First results show its capability to distinguish different process qualities.For this purpose, various configurations for in-process integration of acoustic measurement techniques are discussed and evaluated. The most promising structure-borne sound concept is integrated and tested in a test bed. In a Design of Experiments for specific parameter selection, cubes with different process qualities are produced, and the acoustic signatures are evaluated. For a first prove of concepts a Neuronal Network is trained to classify three different laser classes. Therefore, different NN topologies were tested and the best-found solution had a precision of more than 90%.2Introduction/MotivationAdditive manufacturing (AM) enables a high level of functional integration, the capability of rapid prototyping, individualized products as well as economical production in small series.

As a result, additive manufacturing processes are becoming increasingly important in many industries. The laser beam melting (LBM or more commonly known as SLM) is focused, since it is an established process in the area of prototype construction that stands at the threshold for the employment in large and small series production (Kief, Roschiwal & Schwarz 2015). The SLM process is a powder-bed-based process in which metal powder is melted layer by layer using a laser to form a solid body (Gebhardt 2016).If material defects occur during the process, they lead to an undefined reduction in fatigue resistance and a shortened life of the components (Yadollahi & Shamsaei 2017). This means that components can only be used at highly stressed and safety-relevant points once complex end of line quality assurance has been implemented. Focusing serial production with this manufacturing technique an end of line quality assurance is much too expensive and has to be mitigated. . (Lanza et al.

Focusing serial production with this manufacturing technique an end of line quality assurance is much too expensive and has to be mitigated. . (Lanza et al. 2017)To avoid costly end of line quality assurance, as well as guarantee the durability of the AM parts, a suitable process monitoring method that is integrated into the build process is advantageous. It is essential that the recorded data gets processed in a way that the information helps to ensure the required process quality. With an information about the actual process quality the built process can be stopped, or process parameters be changed to mitigate further defects.Among the different kind of defects inner defects are of special interest. For porosity, some aircraft companies have defined a maximum diameter of 200µm to ensure the durability of AM parts.

For porosity, some aircraft companies have defined a maximum diameter of 200µm to ensure the durability of AM parts. (Kiefel, Scius- Bertrand & Stößel 2018) Other work has shown that the shape and the location in relation to the outer surface has a significant role on the fatigue resistance. (Reschetnik 2017) This makes it necessary to integrate process monitoring techniques that give quantitative information about shape, location and size of inner defects like pores. This work will first investigate current existing techniques upon their capabilities before suggesting anacoustic measuring setup and presenting first results proving this concept.20973 State of the artIn this section a short review of SLM defect characteristics will be given and existing process monitoring techniques are discussed. Furthermore, the approaches to visualize the results of this monitoring techniques are introduced.

Furthermore, the approaches to visualize the results of this monitoring techniques are introduced. This is followed by the discussion about systematically introduced defects in reference test bodies for evaluating the capabilities of process monitoring techniques. The last part of this section focuses on the discussion of relevant reference measurement techniques to evaluate location and shape of pores in AM parts.3.1 Defects in SLMThere are different characteristic defects in SLM (Everton et al. 2016; Grasso & Colosimo 2017). Pores are of special interest since they are difficult to detect by end of line inspection and have a major influence on the fatigue behaviour (Leuders et al. 2013). To evaluate the influence it is important to know the shape and location of each pore (Reschetnik 2017), which strongly correlates with its formation mechanism.Too High EnergyToo Low EnergyKeyhole/VaporizationSpatterUnmolten powder(Gong et al.

2014)(Haeckel 2017) Table 1: Pores may result from either bringing in too high or too low energy in the process zone (Gong et al. 2014)(Gong et al. 2014)When bringing in too high energy keyhole phenomena may occur, which result in subsurface pores. These kind of pores often have a spherical shape and indicate high concentrations of alloys with low vaporization temperature due the high temperature at the bottom of the melt pool, shown in Table 1 (Liu et al. 2016). Another phenomenon which is likely to occur from too high energy input is the generation of spatter which results from the recoil pressure spattering molten material (Khairallah et al. 2016). This molten material is likely to form a spherical shaped drop somewhere on the powder bed. Even though there are a few other phenomena resulting in powder bed contamination in the end this powder bed contamination cannot be melted completely and is likely to trap unmolten powder under it.

Even though there are a few other phenomena resulting in powder bed contamination in the end this powder bed contamination cannot be melted completely and is likely to trap unmolten powder under it. (Haeckel 2017) The micrograph of such pore resulting from powder contamination is shown in Table 1-middle. Not only does spatter lead to insufficiently molten powder, if there is a too small overlap in scan strategies or insufficient laser power input, welding lines are only partly connected like shown in Table 1 on the right. (Gong et al. 2014; Thijs et al. 2010)By tuning the process parameters, it is possible to mitigate pores resulting from too low or too high energy input. Despite this, process and melt pool instabilities from powder bed contamination or attenuation of the laser power from process gases, resulting in discontinuity energy input, can’t be avoided completely. In total there are more than 50 process parameters influencing the SLM process which have to be considered.

In total there are more than 50 process parameters influencing the SLM process which have to be considered. (Spears & Gold 2016; van Elsen 2007) Only few are correctly user controlled and they have only been partially researched. This shows the necessity of process monitoring or process integrated measurement techniques capable of detecting and quantifying pores.3.2 Sensors for process monitoringTo detect process discontinuities which may result in pores, different in-situ process monitoring techniques exist. These are summarized in table 2. They can be divided in optical and acoustic monitoring techniques. Most of these methods are already known from laser welding. (Purtonen, Kalliosaari & Salminen 2014) There are first works also evaluating eddy current testing and backscatter computed tomography. These methods are up to date not yet implemented.

(Purtonen, Kalliosaari & Salminen 2014) There are first works also evaluating eddy current testing and backscatter computed tomography. These methods are up to date not yet implemented. (García-Martín, Gómez-Gil & Vázquez-Sánchez 2011)2098First process monitoring techniques available in industrial application focus on the integration of optical measuring equipment in the beam path (Berumen et al. 2010; Everton et al. 2016). By integrating cameras and diodes/pyrometers into the optical path it is possible to quantify melt pool size, shape and radiation from the melt pool (Clijsters et al. 2014; Furumoto et al. 2013). With this information it is possible to realize a closed-loop control to reduce the occurrence of too high energy input and thereby the number of produced pores can be reduced.

2013). With this information it is possible to realize a closed-loop control to reduce the occurrence of too high energy input and thereby the number of produced pores can be reduced. Further work also showed a correlation of position data of the laser with data from the melt pool monitoring, resulting in maps for each layer which help the operator to identify process variation, but cannot give a quantitative feedback about occurred defects and part quality. (Grasso & Colosimo 2017)A more complex on-axial measurement technique is the optical coherence tomography, where a second laser system is integrated in the beam path (Neef et al. 2014). It might even be upgraded with a separate galvanometer-system scanning the melt pool (Kanko, Sibley & Fraser 2016).

2014). It might even be upgraded with a separate galvanometer-system scanning the melt pool (Kanko, Sibley & Fraser 2016). This technique achieves a new level of information, making it possible to evaluate further melt pool properties like keyhole depth and the geometrical profile of the melt pool.The integration of Off-Axial sensors have the scope to monitor powder bed properties or scan path/scan surface (Grasso & Colosimo 2017). For observing the scan path or scan surface, pyrometers (Doubenskaia, Pavlov & Chivel 2010; Islam et al. 2013) or IR Cameras (Krauss, Eschey & Zaeh 2012) can be integrated. For observing the powder bed, cameras (Zur Jacobsmühlen et al. 2015) or laser line scanners (Haeckel 2017) can be used. For scientific purposes and to understand the process as well as melt pool dynamics, a high speed camera close to the process zone was integrated. (Ly et al.

For scientific purposes and to understand the process as well as melt pool dynamics, a high speed camera close to the process zone was integrated. (Ly et al. 2017)Besides the above discussed optical monitoring methods mentioned, there are novel works elaborating further non-destructive testing techniques. Existing works an all four different kinds of non-destructive testing methods are summarized in Table 2.There are first works on Eddy Current Testing integrated in the process, which right now is related with a high effort scanning the surface and thereby slows down the process. (García-Martín, Gómez-Gil & Vázquez- Sánchez 2011)Rieder integrated an ultrasonic transducer under the build platform which can detect porosity by pulse reflection method. In this case an ultrasonic wave is generated by the transducer and the reflected wave from the backwall or defects is analysed. (Rieder et al. 2016) For this active acoustic approach an acoustic signal must be generated.

(Rieder et al. 2016) For this active acoustic approach an acoustic signal must be generated. The spatially resolved acoustics emissions approach, introduced by (Smith et al. 2016) works in a similar way. Surface acoustic waves are generated with a laser and the propagation is evaluated by measuring the surface acoustic wave at a second point. (Ye et al. 2018) introduced a passive acoustic approach measuring the airborne acoustic signature. Due to plasma generation it is possible to evaluate process discontinuity and detect overheating and balling.Beside of these monitoring techniques first concepts are available for integrating an X-Ray into the process. This approach promises a high degree of information about inner defect. (Shedlock, Edwards & Toh 2011)OpticalEddy CurrentAcousticOff-AxisOn-AxisActivePassive2099CTHigh-speed camera (Ly et al. 2017)IR camera (Krauss, Eschey & Zaeh 2012) Pyrometer (+Camera) (Islam et al.

Not all of them will be used in productive SLM systems in the future. As soon as a technology comes out of the development phase, during which process engineers operate the machine, it is necessary to get intuitive and clear feedback about the process quality. For this purpose, visualization techniques are used to show deviations in a 2D or 3D image like shown in ((Krauss 2016)). Figure 1 shows areas marked in red, in which high IR radiation from the melt pool is observed. This is an indicator for too high energy input and therefore likely to cause defects. (Krauss 2016) This kind of visualizations require interpretation by an expert and may be improved by automated pattern recognition algorithms. (Zenzinger et al. 2014)Besides visualizing process deviations, a classification directly from raw signals is possible. A classification approach using machine learning algorithms for five classes was shown by (Ye et al.

A classification approach using machine learning algorithms for five classes was shown by (Ye et al. 2018), who were able to distinguish balling (a), light balling (b), good process (c), light overheating (d) and overheating (e) like shown in figure 2 by using the air borne acoustic raw signal. This classification of defects in combination with giving an indicator of the severity makes it easier for machine operators or automatized control systems to take the right actions or even precautions to prevent further defects in the process.Figure 1: Data evaluation of a process monitoring technique (Krauss 2016)Figure 2: Distinguishing of classes: balling (a), light balling (b), good process (c), light overheating (d) and overheating (e) (Ye et al. 2018)3.4 Reference bodiesThe systematic introduction of defects or process characteristics in reference bodies is a common approach to evaluate the performance of process monitoring techniques.

2018)3.4 Reference bodiesThe systematic introduction of defects or process characteristics in reference bodies is a common approach to evaluate the performance of process monitoring techniques. The most simple way to do this, is the creation of2100single welding tracks like (Ye et al. 2018). This is especially advantageous since no complex measurement technique is needed to evaluate the properties of the welding tracks.More complex three dimensional reference bodies are used by (Rieder et al. 2016) or (Toeppel et al. 2016). To systematically introduce defects different parameters are used. They range from parameters correlating directly with the energy input (layer thickness, laser power, scan speed, hatching distance) to inert gas flow or powder bed contamination. Table 3 gives an overview over considered defects and the used process parameters.

Table 3 gives an overview over considered defects and the used process parameters. Beside of these used process parameters there are much more process parameters which affect process quality and are not yet used for evaluating process monitoring techniques since they are difficult to vary. (Krauss 2016; Spears & Gold 2016)Research approach (Rieder et al. 2016) (Toeppel et al. 2016)(Ye et al. 2018) (Haeckel 2017) (Zenzinger et al. 2014) (Krauss, Eschey & Zaeh 2012)Considered parameter Density Pores/densityBalling/overheating Pores Pores PoreChanged process parameter Laser power Laser geometry, scanner time delay Laser power Powder bed contamination Inert gas Flow Laser Powerpower,hatchTable 3: Overview of considered defects and varied process parameters3.5 Reference measurement techniqueIf reference bodies are produced it is essential for qualifying the measurement technique to have a reference to measure the introduced defects.

To get information about the inner defects like pores different measurement techniques are suitable, each having its own advantages. In total, three different types are worth to be considered: Archimedes density measurement, micrographs and computer tomography (CT) (Spierings, Schneider & Eggenberger 2011)Density measurements according to Archimedes have the lowest degree of information since it only gives a number of the overall density but no information about the location or particular shape of pores. Beside of this, the reproducibility and standard deviation of Archimedes measurement is the best of the three mentioned measurement techniques. (Spierings, Schneider & Eggenberger 2011)Micrographs give the opportunity to evaluate the shape of pores and the location in the cross-section. This gives more information about pore properties than Archimedes measurement. By using microscopes, the usable resolution of this approach goes down to a few micrometres.

This gives more information about pore properties than Archimedes measurement. By using microscopes, the usable resolution of this approach goes down to a few micrometres. Unfortunately the process of making micrographs with grinding, polishing and the parameters for taking pictures to evaluate the pores are hard to reproduce. (Spierings, Schneider & Eggenberger 2011)The highest amount of data about the reference body can be generated by using a CT. With CT it is possible to evaluate size and shape of the pores according to BDG P202 (BDG 2010) over the whole volume of the test body. Major drawback of this technique is its measurement uncertainty and its undefined limit for small pores. Depending on the used parameter for the image acquisition the resolution of the CT image changes. (Schild et al. 2018)2101distance,4 Methodical approachIn this section the overall approach for developing an acoustic monitoring system will be explaind.

(Schild et al. 2018)2101distance,4 Methodical approachIn this section the overall approach for developing an acoustic monitoring system will be explaind. Therefor first of all the reasons for investigating an acoustic measurement setup will be explained before its integration into a test bench will be shown. In section 4.2 de design of the chosen reference bodies and the chosen defect mechanisms are introduced. Section 4.3 focuses on the data processing of the acoustic signal and the correlation of the features generated from the reference measurement.4.1 Sensor and test setup4.1.1 Selection of measurement techniqueFigure 3 summarizes the capabilities of the different non-destructive testing (NDT) methods in general and compares their abilities in resolution and defect position (inside or surface) in a qualitative way. Since eddy current testing is only capable of detecting defects close to surface all techniques above the red line can detect defect inside a volume.

Since eddy current testing is only capable of detecting defects close to surface all techniques above the red line can detect defect inside a volume. The ones below the red line can detect surface or close to surface defects.Figure 3: Comparison of NDT methods in resolution and defect positioning according to Gevatter & Grünhaupt 2006)Section 3.1 showed that the formation mechanism of SLM characteristic pores only partly understood and there are works indicating the pore formation taking place without being visible at the surface (Calta et al. 2018). Nevertheless, pores are volume defects which makes it necessary to have a closer look on the testing methods which are in general capable of evaluating defects within the volume.X-ray backscatter tomography, which is a special method of the computed tomography (CT), is part of further research and not yet integrated in the process. The process integration of CT was already shown for FDM and is also plausible for SLM.

The process integration of CT was already shown for FDM and is also plausible for SLM. Regarding the high integration effort and investment cost, this is likely to be limited for research and special applications. Since the integration of acoustic sensors is much less complex (as shown in (Rieder et al. 2016; Ye et al. 2018)), it is promising to have a closer look on this kind of measuring principle.4.1.2 Integration concepts of acoustic measurement systemsThe generation and detection of acoustic signals in general can be performed by three different kind of actors/sensors: piezo, electromagnetic transducer (EMAT) or by the laser itself. Furthermore, acoustic waves can be air borne or structure born. This work will focus on structure born noise due to its expected higher degree of information.With regard to the integration of the actor and the sensors, there are two general placements, which are schematically sketched in figure 5 as variants A and B.

In variant B, the probe is mounted in the available space between the construction platform and the substrate plate. Such a construction has already been implemented and tested (Rieder et al. 2016). Variant A provides space for the sensor to be positioned above the workpiece or 2102integration into the beam. The installation space above the workpiece is limited by the fact that the beam path of the laser must not be disturbed. Nevertheless, this variant offers advantages in terms of component coverage.Figure 4: Integration and signal generation/detection options for SLM process monitoringWith these options to integrate sensors and the three sensor options to generate/detect acoustic signals, 18 different options for acoustic testing in SLM can be generated. 9 concepts are plausible and worth of further discussion which can be found in (Eschner et al.

9 concepts are plausible and worth of further discussion which can be found in (Eschner et al. 2017) and will be only summarized here.In general, the integration above in volume A of a piezo electric or EMAT makes it necessary to implement a mechanism to move the transducer for signal detection over the process zone in order to not interact with the laser beam. Such mechanisms are complex, susceptible to failure and slow down the overall process time.When the laser induces energy, acoustic waves are generated which can be measured and evaluated like((Ye et al. 2018)) has shown for air-borne acoustic emissions.

2018)) has shown for air-borne acoustic emissions. It is likely to interact with ferromagnetic material and EMATs which lead to a disturbance of the powder bed.Low frequency ultrasonic transducers have limited performance in resolution, while high frequency ultrasonic transducers have limited access to the whole volume since the waves are damped excessivelyFor the integration in variant A attention should be spent to the existing temperatures and the platform heating. Nevertheless, it was shown, that the integration at this position is possible.To have an optimal solution regarding process time, expected information content and minimizing integration effort of the nine plausible concept one was favoured.

The favoured concept is shown in figure 6 and will be explained in the following.2103Figure 5: Monitoring concept for SLM using structure-borne acoustic emissionsIn this concept, the acoustic waves are generated by the SLM process itself from plasma and other process dynamics. The generated waves propagate through the build object and are recorded by a piezoelectric transducer. If a defect, like a crack prevents the propagation recorded signals change. Furthermore, the process stability itself can be judged, since process inhomogeneity produces inhomogeneous signals. This concept is favoured since it monitors the building process and thereby promises the highest degree of information, while not slowing down the process. (Eschner et al.

This concept is favoured since it monitors the building process and thereby promises the highest degree of information, while not slowing down the process. (Eschner et al. 2017)4.1.3 Integration of an acoustic measurement system in a reference SLM processIn order to investigate the capabilities of the process monitoring technology, reference bodies are printed as shown in section 3.4 and systematic errors are introduced into them. For this purpose, it is necessary to have a process environment which gives the flexibility to vary process parameters freely and integrate the sensor. A test setup imitating the SLM process was designed, giving flexibility for the relevant process parameters and sensor integration.Figure 6: Sectional view of SLM build chamber in test setupFigure 7: Complete view of test setup2104Figure 6 shows a cross sectional view of the process chamber which is the central unit of the test setup.

The powder application mechanism is a simple wiper mechanism to reduce overall complexity. There is one cylinder used as a powder reservoir. The wiper moves from left to right and applies the powder from the powder reservoir onto the build platform. The build platform moves down after a layer is exposed. Under the build platform an acoustic sensor by QASS is integrated. All axes are moved with stepper motors in micro stepping mode. Right next to the build platform, outlet and inlets for the inert gas flow are integrated. The outlets and inlets are easily replaced to influence the inert gas flow. The whole system is sealed against the environment to avoid oxygen contamination and to realize positive or negative pressure. The process chamber is integrated in a machine frame made from standard profiles which connect all the different components (Figure 7).

The process chamber is integrated in a machine frame made from standard profiles which connect all the different components (Figure 7). Since an acoustic sensor is used, some effort was spent for acoustic decoupling of peripheral equipment like computer or optical system with the scan head.The laser system consists of a 250W IPG laser and a Cambridge Technology SMC scanner card. The XML API provided by Cambridge Technology offers a flexible adaption of process parameters. A python script generates the code needed for every layer of the test cubes. The cubes are then placed in a 3x3 matrix on the substrate plate. Every cube is printed with its own set of parameters. More complex objects are generated via a CAD/CAM chain consisting of a slicer software (Slic3r with a custom configuration suitable for SLM) and a python script converting the generated G-code into the according xml code. The printing parameters are also freely variable.

The printing parameters are also freely variable. A main LabVIEW application controls the stepper motors, the inert gas system, all sensors as well as the QASS ultrasonic measurement device. The data is recorded with a frequency of 4 MHz. The inert gas system is designed to have the ability to vary process parameters like flow rate and oxygen content. For this purpose, the system is equipped with an oxygen-, a temperature- and a flowrate sensor.To get information about the actual position of the laser beam spot and to synchronize this information with the acoustic signal, a FPGA for interpreting the XY2-100 protocol (protocol for communication between scanner card and galvanometer scanner) of the galvanometer system was realized.

With this system it is possible to assign a position to its corresponding acoustic signal during the build job.4.2 Evaluation of the acoustic measuring system with reference bodiesTo evaluate the acoustic measuring system in a first approach, reference bodies are printed with varying part quality. As shown in section 3.4 there are different ways for producing reference bodies. One of the easiest ways is to vary process parameters like laser power ( ) or ), layer thickness ( hatching distance ( ). These four parameters directly correlate with the energy density which can be calculated as following:), scan speed (𝑃𝑃𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑑𝑑𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑣𝑣𝑙𝑙𝑠𝑠𝑙𝑙𝑠𝑠𝑑𝑑ℎ𝑙𝑙𝑎𝑎𝑠𝑠ℎ𝑖𝑖𝑠𝑠𝑖𝑖(Formula 1)𝑃𝑃𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙 𝑣𝑣𝑙𝑙𝑠𝑠𝑙𝑙𝑠𝑠 ∗ 𝑑𝑑𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙 ∗ 𝑑𝑑ℎ𝑙𝑙𝑎𝑎𝑠𝑠ℎ𝑖𝑖𝑠𝑠𝑖𝑖𝐸𝐸 =Previous work by (Cherry et al. 2015) (see Figure 8) and (Gong et al. 2014) (see Figure 9) have shown that the energy density within the process is correlated directly to the part quality, which was measured in these cases on porosity.

2014) (see Figure 9) have shown that the energy density within the process is correlated directly to the part quality, which was measured in these cases on porosity. There is an optimal interval, in which the process produces good results. Above and below the interval, quality drops rapidly (Gong et al. 2014; Cherry et al. 2015). Based on this knowledge, the varied process parameter to influence part quality regarding pores will be the energy density used in the process.2105Part Porosity vs. Laser Input Energy Density25Part Porosity vs. Energy Density1020]]%[ y t i s o r o P864top%[ y t i s o r o P151052side00501001500050100150200250Energy density [10E9 J/m^2]Laser input energy density [J/mm3]160W 120W 80W 40WFigure 8: Relation between part porosity and energy density in (Cherry et al. 2015)Figure 9: Relation between part porosity and energy density in (Gong et al.

2015)Figure 9: Relation between part porosity and energy density in (Gong et al. 2014)To evaluate the capabilities of the acoustic monitoring technique, this correlation of energy density and part quality/porosity is used. Therefore, to obtain a wide range of energy inputs, a Design of Experiments (DOE) was conducted with the four parameters influencing the energy input. Since only four parameters are varied, a full factorial test plan is used. Each parameter range has a minimal value, a centre point and a maximal value.The given formula for energy input (Formula 1) is a hyperbola, which means a true centre point of the min-max interval results in an unevenly distributed energy input throughout the test objects. To counter this effect, the centre value of the scan speed was set to a third of the interval range.

To counter this effect, the centre value of the scan speed was set to a third of the interval range. The parameters were set as stated in Table 4:Parameter Laser power Scan speed Layer thickness Hatching distanceMinimal value 150 W 200 mm/s 0.02 mm 0.03 mmCentre value 200 W 400 mm/s 0.03 mm 0.04 mmMaximal value 250 W 1000 mm/s 0.04 mm 0.05 mmTable 4: Parameter set for Design of ExperimentsThe resulting 81 parameter sets are printed in the designed test setup as cubes with an edge length of 0.5 mm. The material used is 316L stainless steel. Each build job contains 9 different reference bodies as shown in Figure 10. Since a 3x3 matrix is printed on the substrate plate, the scan speed is varied in the y direction, the laser power in x direction (compare Figure 10). The other two parameters, layer thickness and hatching distance, are varied in between the different substrate plates. This results in a total of 81 test cubes on 9 substrate plates.

The other two parameters, layer thickness and hatching distance, are varied in between the different substrate plates. This results in a total of 81 test cubes on 9 substrate plates. To ensure an even distribution over all energy input levels, the energy input over the parameters as well as the number of the according substrate plate can be plotted (Figure 11)2106200Figure 10: Substrate plate with 3x3 cube matrix and varied print parametersDistribution of Energy Densities in Substrate Plates2500.002000.00] 3 ^ m m/ J [ y t i s n e d y g r e n E1500.001000.00500.000.00012345678910Plate numberFigure 11: Energy density distribution over all substrate platesSeveral tests conducted before setting up the test plan showed a rough estimate of the expected process window. An energy input between 500 and 1000 seems to be optimal.

An energy input between 500 and 1000 seems to be optimal. For testing purposes, it is mandatory to achieve values below and under that interval, thus provoking process states with too high energy (resulting in keyhole mode / overheating) and states with too low energy (pores, partial melting of powder).3𝐽𝐽/𝑚𝑚𝑚𝑚2107Platte 1Platte 2Platte 3Platte 4Platte 5Platte 6Platte 7Platte 8Platte 94.3 Correlating acoustic signals with reference measurementsThe main objective of the utilised process monitoring technique is not to detect the variance between different processes, but rather to carry out an analysis of the component quality (with the associated defects) at coordinate level. The goal is to automatically determine the type of defect as well as the spatial extent and location in the component. Selected process parameters are to be validated automatically by a machine learning technique.

Selected process parameters are to be validated automatically by a machine learning technique. If it is necessary to adjust parameters due to the current process situation, the system should propose a suitable parameter set.To evaluate the printed test cubes, two methods will be used: computer tomography and Archimedes. The results of the reference measuring, the acoustic signal of the process (with extracted features) and the printing parameters are then compiled into datasets. These are then used to train an artificial neural network (ANN), which will be realized in Pythons sklearn package and Google TensorFlow. To handle the great amount of data, most calculations will be done with GPU support (Nvidia CUDA).For first analysis classification is used to determine the type of the main defect (overheating, balling, pores), whereas in future work regression will be used to get quantitative features like density value.

The complete dataset is randomly split into training (70%) and test data (30%) to test repeatability of the training (each training is repeated five times and the according mean and standard deviation is calculated). To evaluate the model performce regarding a classification over the whole dataset, a five-time cross-validation is performed with an according 80/20-split of the data.The raw input signal of the sensor is decoded from its binary and proprietary data format of the manufacturer into a Pandas dataframe which is stored with zlib compression as a HDF5 file on disk. The timeseries in then fast fourier transformed (FFT) using a window of 0.005 seconds. As windowing function, the hanning window is used. To reduce background noise, a frequency mask of the background signal is subtracted from the resulting FFT (see Figure 13 and Figure 14).

As windowing function, the hanning window is used. To reduce background noise, a frequency mask of the background signal is subtracted from the resulting FFT (see Figure 13 and Figure 14). The result is a spectrogram for each layer like shown in Figure 12.Figure 12: Spectrogram of measured signal of one layer (filtered)To feed the FFT of each layer into the artificial neural network, the data must be flattened since they are stored as an array (Formula 2). The additional features can be added in the future to the input vector. The vector resulting directly from the spectrogram contains 12 million values, meaning one value for each pixel. To reduce2108this amount, a bottleneck approach is used. The number of neurons in a hidden layer is set between five and a hundred.

The results contain the analysis of 18 cubes build in two build jobs. Each cube has 120 usable layers (the first five layers were discarded due to double exposure), which leads to a total dataset of 2160 individual datasets. To gain first experiences with artificial intelligence (AI) data analysis, only the given laser power from the output vector was used. So, no features from the reference measurement technique is used for first exemplary evaluations of acoustic signals of the process. The analysis of all 81 cubes with the use of a reference measurement technique will take place in the future.5.1 Capturing raw process dataFor data processing, it is crucial to obtain a clear, noise free signal with a sufficient amplitude. First tests with a ceramic coupling between the sensor and the substrate plate showed a high dampening of the process sound. Therefore, the resulting measurements had to be discarded.

First tests with a ceramic coupling between the sensor and the substrate plate showed a high dampening of the process sound. Therefore, the resulting measurements had to be discarded. After removing the ceramic coupling and adding glycerol as a coupling agent, the signal was much stronger. As a drawback, the background noise of the test setup (stepper motors, vibrations of the inert gas pump) were present and reduced the usable information within the signal. Using a difference mask (see Figure 13 and Figure 14) showed very good results in reducing noise from the test setup. The printing process of each layer was recorded by the QASS measurement system. The resulting time series was then transformed via FFT to the frequency domain.

The printing process of each layer was recorded by the QASS measurement system. The resulting time series was then transformed via FFT to the frequency domain. The process has an effective frequency range up to 1.3 MHz.Figure 13: Raw measuring signal without applied filterFigure 14: Difference mask to reduce noise from test setup5.2 Data preprocessingEach measurement is decoded and processed and stored along with its parameter set on disk. One substrate plate creates roughly 100 gigabytes of data. Every layer is then assigned with an output class: for the first test, the used laser power. For feeding the data into an artificial neural network, every dataset must contain an equal number of samples. To achieve this, the machine noise is appended at the end to fill to the desired length. During pre-processing, a dataset with input and output vectors is created and stored separately on disk to reduce processor (CPU) time for repeated use.

During pre-processing, a dataset with input and output vectors is created and stored separately on disk to reduce processor (CPU) time for repeated use. The dataset is then randomised and split into 70% training data und 30% test data.5.3 Artificial IntelligenceUtilizing sklearn’s MLPClassifier1, a multi-layer perceptron is used to make a first attempt on classification. First results show a better performance of a stochastic gradient descent solver over the ADAM solver. As activation function, ReLu is used.

First results show a better performance of a stochastic gradient descent solver over the ADAM solver. As activation function, ReLu is used. Tanh was tested but produced worse results than ReLu.1 Scikitlearn: A python package which has a lot of built in AI and data processing functionality2110Precision defines the number of true positive classifications in regard to all positive classifications (“What proportion of positive identifications was actually correct?”2), whereas recall is calculated by comparing the true negative classifications to all real positives (“What proportion of actual positives was identified correctly?”3). The F1-Score combines precision and recall via (Formula 4):(Formula 4)2 ∗ (𝑜𝑜𝑝𝑝𝑐𝑐𝑐𝑐𝑙𝑙𝑙𝑙𝑙𝑙𝑜𝑜𝑝𝑝 ∗ 𝑝𝑝𝑐𝑐𝑐𝑐𝑐𝑐𝑙𝑙𝑙𝑙) 𝑜𝑜𝑝𝑝𝑐𝑐𝑐𝑐𝑙𝑙𝑙𝑙𝑙𝑙𝑜𝑜𝑝𝑝 + 𝑝𝑝𝑐𝑐𝑐𝑐𝑐𝑐𝑙𝑙𝑙𝑙𝐹𝐹1 =To check the repeatability of training the artificial neural network, the network was trained and tested five times on the same split of training and test data.

The mean and standard deviation of precision, recall and F1- score are shown in Figure 15 to Figure 20.When regarding precision, recall and f1-score, one can notice that there is an interval of neurons in which the classification results are optimal. Above and below that interval the classification score drops rapidly. Furthermore, it can be seen that the addition of multiple hidden layers does not provide a increase regarding the reachable classification score. In contrary, when regarding the standard deviation it is noticeable that the standard deviation rises with the addition of more layers.

The results of the cross validation are shown 2112in Table 5. The selected topology does not have quite as strong an effect through complete cross-validation. The achievable classification rates between 80 and 90 percent represent a good level for a first approach (no tweaking regarding learning rate, momentum etc. was done).Model (neurons in one hidden layer)MeanPrecisionStandard deviationMeanRecallStandard deviationMeanF1-ScoreStandard deviation20 30 40 500.85 0.86 0.87 0.870.026 0.024 0.027 0.0380.83 0.86 0.87 0.870.031 0.025 0.027 0.0410.83 0.86 0.87 0.870.031 0.025 0.027 0.038Table 5: Results of a fivefold cross-validationA major drawback when working with such big datasets is computation performance. Training duration rises linear with the number of neurons in a hidden layer until a certain point, where it starts rising exponentially (Figure 21).

Training duration rises linear with the number of neurons in a hidden layer until a certain point, where it starts rising exponentially (Figure 21). This is probably due to the internal memory of the computer maxing out, so the operating system starts to cache data on the system disk. This process greatly reduces the speed of the training. Hence it is very important to have enough internal memory to not slow down the process. Furthermore, the batch size should not be too small. The overhead from various function calls increase the training time as well (Figure 22). Last, the final model will be stored on disk. Using a single hidden layer with a hundred neurons, the final model reaches almost 20 gigabytes (Figure 23).

For further use, the model is loaded in memory, which decreases the available memory for batches or limits the maximum number neurons trainable.] h [ e m TiTraining duration (batch size constant)1.61.41.21.00.80.60.40.20.050 15 10 25 20 30 40 55 70 0 1Number of neurons in hidden layerTraining duration (neurons per hidden layer constant)] n mi[ e m Ti45 40 35 30 25 20 15 10 5 00 10 30 50 70 90 1 1Batch size0 3 10 5 1] e t y B G[ e z i s el iFTrained model size20 18 16 14 12 10 8 6 4 2 050 15 10 25 20 30 40 55 70 0 1Number of neurons in hidden layerFigure 21: Training duration changes over hidden layer sizeFigure 22: Training duration decreases with increasing batch sizeFigure 23: Final models get bigger with increasing complexityRegarding the current classification the following statements can be made:Bottleneck approach is very useful, since roughly 12 million input neurons are used • Three hidden-layer-topology offers best classification result, but also needs more computational resourcesOptimum between 25 and 45 neurons in hidden layer (also valid for multi-hidden-layer topologies) • Maximum precision with this approach: 0.93 • Maximum recall with this approach: 0.93 • Maximum F1-score with this approach: 0.93 • ADAM is slower than SGD 2113These findings give a first impression how the artificial neural networks can be suitably designed toevaluate the acoustic signals of the process for an in-situ process monitoring.6 Summary and OutlookTo detect pores in additive manufactured parts, a possible solution lies within an in-situ process monitoring.

As a measurement technique, an acoustic sensor was chosen and implemented. The acquired process signals could be processed and were fed into an artificial neural network. It was possible to make a classification regarding the used laser power in the process. Multiple topologies of the artificial neural network were tested. A bottleneck approach with a single hidden layer seems the most promising. It can be concluded, that the acoustic signal of the process is suited to characterize the SLM process.Once the final DOE is finished (all plates are printed), a bigger dataset will be created, and the different models will be tested against varying output classes. Using Google TensorFlow might decrease the overall time needed for training and testing.

Using Google TensorFlow might decrease the overall time needed for training and testing. Furthermore, different parameter sets must be tested as well as other types of neural networks like convolutional neural networks or restricted Boltzmann machines.After getting reference measures of the defects in the parts, it will be tried to correlate the defect type and position to the acoustic signals. The AI approach can easily be supplemented with additional input and output parameters to enable modular traceability.

VELOPMENT OF AN ADDITIVE LAYER MANUFACTURING (ALM) SELECTION TOOL FOR DIRECT MANUFACTURE OF PRODUCTSP. C. Smith and A. E. W. Rennie paul.smith@lancaster.ac.uk and a.rennie@lancaster.ac.uk Lancaster Product Development Unit, Lancaster University, Lancaster, LA1 4YR, UK Reviewed, accepted September 10, 2008ABSTRACTAdvancements in Additive Layer Manufacturing (ALM) technology and a broader knowledge base of process and material capabilities make ALM increasingly, a more valid manufacturing option.Small creative industry and industrial designers in the UK, as well as experienced engineers, can benefit from the freedom from design, manufacture and distribution constraints that ALM technology offers, yet they are unaware of the opportunities available.This paper present a method for selecting an ALM technology as a manufacturing method, based on a part specification, as an ALM selection tool.

Selecting appropriate processes, materials and giving design for rapid manufacture advice are part of the recommendations offered from this ALM selection tool.KEYWORDS Additive Layer Manufacturing, Design for Rapid Manufacture, Selection Tool, Creative IndustryINTRODUCTIONRapid Manufacturing (RM) using ALM technologies, also known as tool less manufacturing, has the potential to allow small companies to manufacture products on very low scales, possible at a production scale of one, thus opening up the opportunity to compete with bigger companies for whom amortising the cost of a tool over a large production run is not problematic.RM with ALM technologies can produce high value products, where value is added through the design of the product, either through innovation or customisation.

Other advantages of RM with ALM technologies are that designers can adopt new design methodologies to improve their products, such as design for assembly and design for service (Ulrich & Eppinger, 2006).Adding functionality to designs to aid assembly, maintenance and disassembly at the end of the products lifecycle theoretically adds little extra cost to its manufacture using ALM. In fact, the potential for radically different design methodologies is one of the major drivers for the development of rapid manufacturing systems and materials (Hopkinson et al. 2006)However, lack of knowledge concerning ALM technologies is slowing the adoption of RM into current design methodologies and so is having little impact on design cycles and is yet to impart any change in the small business sector in the UK for design, despite its potential for change. So far, in terms of RM, we do not yet have a full understanding of all the features that may be required for a typical part (Hopkinson et al.

So far, in terms of RM, we do not yet have a full understanding of all the features that may be required for a typical part (Hopkinson et al. 2006). On top of the lack of understanding of RM parts, the market is flooded with technology and materials that are not yet fully understood and so selection of suitable technologies is a laborious task even for expert users of ALM. Forsmall businesses and non-experts, the task can be impossible, so RM using ALM remains underutilised.Cost of RM using ALM is another inhibiting factor for its adoption. Yet, in order for the cost of ALM technologies to come down, there needs to be an increase in their use, and an increase in demand can drive the growth of the ALM market bringing in new machine and material options for designers to utilise. Conversely, in order for companies, especially small to medium sized companies, to start exploiting the advantages of ALM, they first need to be made aware of what is possible now and what is available.

This paper presents a potential solution to these problems as an ALM selection advisor tool.The research presented in this paper is a review of previous research regarding selection of an ALM technology as a method for producing a rapid prototype as opposed to selection of a RM method. The paper then describes the building of a relationship database of available ALM technologies based on what is currently available in the UK market through service providers and bureaux. The paper then goes on to present a tool which can be queried by a user to recommend suitable technologies to manufacture any given product. The final part of the paper describes further work to be undertaken.EXISTING SELECTION TOOLSIntelligent RP System SelectorThe Industrial Research Institute Swinburne (IRIS), based at Swinburne University in Australia, was developing a system named the Intelligent RP System Selector in 2001.

It is an expert rule based system that assists a user wanting to purchase a rapid prototyping technology. This is done by asking the user a series of questions relating to their machine specification. The main criteria for selection are price, dimensional accuracy, surface finish, maximum build volume, range of materials, range of layer thickness and speed of build (Masood & Soo, 2002)The system is aimed at both experienced and novice rapid prototype users in manufacturing. It includes rapid prototype systems from the USA, Japan, Germany, and Israel and includes 39 rapid prototype machines available before 1998. The selection criterion for the system was based on results of a questionnaire that was given to groups of vendors and users.The program has four main choices for the user to make: Quick Selection, Detailed Selection, Build Technology, and Machine Style. An illustration of the system user interface (UI) is given below in Figure 1.

An illustration of the system user interface (UI) is given below in Figure 1. The selector is programmed with a series of ‘IF’ and ‘THEN’ rules based on the knowledge input into the system. The system checks user input against the rules and generates recommendations based on that input.Figure 1: Intelligent RP Selector User Interface (UI)Data for the selector system was gathered from original equipment manufacturers and verified through vendors and users of the technology. For each of the four major selection options, a different basis is used to generate the recommendation:Quick Selection: Price, XY accuracy, build volume and material.

For each of the four major selection options, a different basis is used to generate the recommendation:Quick Selection: Price, XY accuracy, build volume and material. • Detailed Selection: Price, XY accuracy, Z accuracy, surface finish, build volume, build thickness and build speed.Building Technology: Laser or non-laser system, building process, price, XY accuracy and build volume.Machine Style: Office environment, desktop, normal commercial type, price, XY accuracy and build volume.The selector asks questions of the user giving them options to choose from. The process continues and an appropriate technology is recommended. Along with the technology description, other information is displayed including vendor details.

The process continues and an appropriate technology is recommended. Along with the technology description, other information is displayed including vendor details. Multiple recommendations can be made by the selector and displayed as a list for the user to choose from.The creators of the IRIS rapid prototype selector suggest that improvements could be madewith the addition of images of the machines and graphs displaying any technical data.Computer based rapid prototyping design advice systemIn 1999, the Design Engineering Research Centre at the University of Wales Institute (Cardiff) developed a computer based rapid prototyping design advice system. The system operates on asimilar basis to the IRIS system which works using a knowledge base and some sort of inference engine, using rules often expressed as ‘IF’, ‘THEN’ and ‘ELSE’ statements organised as a pattern.

However, this system works by breaking tasks down into subroutines, some of which make decisions based on input data and others contain rules for checking and either implementing or rejecting the decision (Bibb et al, 1999). This system gives priority to certain input criteria and other input is checked against this initial selection.This particular system also incorporated the use of Computer Aided Design (CAD) data in the selection process, primarily because the overall design of the system was to be user friendly and so user input was kept to a minimum. The CAD data was interpreted and values inferred from it through calculations from a second program. The resulting values were then used as input data for the main system.

The CAD data was interpreted and values inferred from it through calculations from a second program. The resulting values were then used as input data for the main system. Having the ability to call secondary programs from the main program allowed for less time to be spent coding the system and more time to be spent on the content (Bibb et al, 1999).Rapid prototype process selection using graph theory and matrix approachRao and Padmanabhan of the National Institute of Technology, Gujarat, India and the SNG College of Engineering, Kerala, India, discussed rapid prototype process selection using graph theory and matrix approach (Rao & Padmanabhan, 2007).Their logical system, like other technology selectors, defines desirable attributes of a rapid prototyping system as process selection criteria.

The interrelations between the selection criteria in terms of their relative importance are modelled in a diagraph and matrix style.The selection criteria they impose are similar to other rapid prototype selectors and include dimensional accuracy, surface roughness, tensile strength, elongation, part cost and build time. In the system, all attributes with the exception of part cost and build time, are given real values. Part cost and build time are expressed as a fuzzy value, for example, very high, medium, low, very low. These fuzzy values are quantified through a conversion table that assign an attribute: a value based on a relative scale that starts at exceptionally low (0.045) and ends with exceptionally high (0.995).A user then defines what the most important attributes are that they require for a model produced using a rapid prototype technology.

The system calculates a value for each technology in the program in relation to the attributes defined by the user, ranking the technologies in order of suitability for the users’ purpose and thus making a recommendation.Rapid prototype system selection system using a TOPSIS methodByun and Lee (2005) describe a rapid prototype technology selection system using a modified technique of order preference by a similarity to ideal solution (TOPSIS) method. Like in other systems, the authors define attributes a user will input to determine the suitability of a rapid prototype technology for production of an end use model. They use both crisp and fuzzy data related to the defining attributes to rank individual rapid prototype technologies. The crisp data they use includes values for accuracy, surface roughness, tensile strength and elongation.

The crisp data they use includes values for accuracy, surface roughness, tensile strength and elongation. These values were obtained via benchmarking of each rapid prototype technology by designing a test part to be produced using each of the rapid prototype technologies. Values for the crisp dataattributes were taken from these test parts. Fuzzy data which includes model price and build time, have values determined from linguistic terms such as very very low to very very high.The attributes used to generate recommendations for a rapid prototype technology for a user are the same as those used by Rao and Padmanabhan in their selector system using graph theory and matrix approach.

In fact, Rao and Padmanabhan borrowed theses attributes from work carried out by Byun and Lee, who determined these attributes through questionnaires with vendors, institutions and bureaus.Once the rapid prototype technologies have been ranked according to their attributes, the user can generate recommendations on which system to choose by giving weighting factors to the various attributes in terms of the most desirable for a particular model. This presents an easy to use system which can handle both crisp and non crisp data. However, no UI has been presented by the authors where it is offered as a system where data can be organised and analysed to generate an outcome.Summary One of the problems with the aforementioned selector tool concepts is that the user needs to have some specific input data about their product before they can make any recommendations using the system. For some designers with very detailed design specifications, this would not be an inhibiting factor.

For some designers with very detailed design specifications, this would not be an inhibiting factor. However, for an industrial designer or small creative industry business, the input data required as logical values such as surface roughness, could be off-putting and to omit it could cause the system to fail. For a system utilised for selecting an ALM technology for rapid manufacturing, the input data needs to be easily understood by the majority of designers, such as using text to describe attributes a product might need as apposed to a numerical value.SELECTION SYSTEM DATABASEIn order for a system to be developed to aid designers in choosing an ALM technology route for manufacturing, firstly a database needed to be built containing all current rapid manufacturing technologies.

For this selector, a relationship database was constructed detailing the types of machines available from various producers, the different types of technologies each machine used, the available materials for each machine and technology, and the characteristics of a part created with a combination of machine and material. Figure 2 below illustrates the construction of the database and the relationships between each category. Constructing the database this way makes it easier later to programme a selector tool to make recommendations based on user input. Each relationship shown here, such as the relationship between one manufacturer who can produce many products, is a database of information relating to that particular category. Figure 3 below illustrates the materials database.

Figure 3 below illustrates the materials database. Each material in Figure 3 has an identification number so that it can be referenced quickly and easily throughout the other relevant databases, and a category identification which puts it into a material category of either polymer, metal or ceramic. The categorisation of attributes in the system, which includes material type, technology type, manufacturer, products, etc, means that building of relationship databases can be done quickly and easily. It also allows the system to be easily queried to generate recommendations and allows it to be easily updated.Figure 2: Relationship database construction for RM selectorThe database contains 40 ALM technologies/machines from seven different manufacturers(see Figure 4 below).

Although there are many more manufacturers of ALM technologies available, the technologies included in this database are based on information gathered by the authors in relation to the most suitable technologies for rapid manufacture of parts and the availability of technologies in the UK market.Figure 3: Materials database part of RM selector databaseFigure 4: ALM technologies contained in the databaseUK AVAILABILITYThe content for the database is related to what is available currently in the UK market through service providers. This is because the RM selector tool which uses the database is aimed at UK creative industry businesses and so it was deemed appropriate to design the system in the context of what was available in the UK.Figure 5 illustrates the availability of technologies in the UK’s service providers and universities.

Figure 6 shows how Stereolithography (SLA) and Selective Laser Sintering are the dominant technology platforms and most widely available to rapidly manufacture products, with a heavy weighting of polymer based systems over metallic systems (Figure 7).The same can be said for the availability of materials for RM in the UK, with SLA resins dominating the material availability (Figure 8).RAPID MANUFACTURING SELECTION TOOL Whereas previous attempts at a rapid prototype selection system have required some level of user input knowledge about either ALM technologies or specific product data, like surface roughness, build orientation etc., this selector uses input data that does not discriminate against lack of knowledge.

It was decided that initially, the use of keywords relating to the required characteristics of any given technology would be used to search the database and generate a list of recommendations that the user could then use to make their final decision on a suitable ALM(cid:3)technology for manufacture. On a user entering certain keywords into the selector tool, it would search for materials that concur to the entered data until it finds a match, or in most cases, many matches. The construction of the relationship databases allows the system to relate each material to a technology, manufacturer and product which will stand as the information generated for the user to view.Figure 6: Percentage of technology platforms in the UKFigure 7: Percentage of polymer systems to metallic systemsFigure 9 below shows the UI for the RM selection advisor programme. It is a web based programme that has been designed to be easy to use.

It is a web based programme that has been designed to be easy to use. Once the user has entered a keyword intothe selector advisor or answered some product profiling questions, the programme searches the database for any relating keywords or materials that match the profile of the product.Figure 8: Availability of materials in the UK by percentageFigure 9: RM selection advisor UIWhen the search is complete, the user is presented with a number of suitable RM materials that relate to a particular technology which can then be used to specify a rapid manufacturing process for manufacture of their product. Included with the recommendations are a series of technical data sheets that relate to each of the materials featured in the database. These sheets are included as support for designers. They give information about each particular material, its corresponding technology platform and advice on how to redesign parts specifically for the recommended material and process.

They give information about each particular material, its corresponding technology platform and advice on how to redesign parts specifically for the recommended material and process. The sheet is available as a PDF document which can be accessed by clicking the document link next to the corresponding recommendation (Figure 10).Figure 10: Recommendations made by the tool and corresponding technical documentThe resulting information is intended to be enough so that a designer could research suitable service bureaus and use it to specify to the bureau what they require. It is not the intention of the selector tool to give definitive information on one recommendation, mainly because there is, for most products, more than one rapid manufactured solution, but also because what is achievable by one bureau service provider with a certain material and technology may not be achievable with another due to experience or conditions.

For these reasons consulting with a service provider would be necessary before a final decision is made.CONCLUSIONS AND FURTHER WORKTo increase the use of ALM technologies as a direct manufacturing method, relevant information needs to be made available to designers of all abilities. The selector tool presented in this paper allows basic specifications to be made for individual products which can then be used to provide recommendations of suitable materials and technologies that can manufacture that product.There are improvements that could be made to the selector tool. As well as improving the overall aesthetic view of the UI, the intention is to include approximate weight or cost values for any given product, the data of which will be input into the system in a future version, and allow the user to select a material category before querying the system.

VELOPMENT OF AN AUTOMATED MULTIPLE MATERIAL STEREOLITHOGRAPHY MACHINEAsim Inamdar, Marco Magana, Frank Medina, Yinko Grajeda, Ryan Wicker University of Texas at El Paso, W.M. Keck Border Biomedical Manufacturing and Engineering Laboratory, Department of Mechanical Engineering, El Paso, Texas 79968-0521AbstractAn automated Multiple Material Stereolithography (MMSL) machine was developed by integrating components of a 3D Systems 250/50 stereolithography (SL) machine in a separate stand-alone system and adapting them to function with additional components required for MMSL operation. We previously reported retrofitting a 250/50 SL machine with multiple vats to accommodate multiple material fabrication for building a wide variety of multi-material models (Wicker et al., 2004). In the MMSL retrofit, spatial constraints limited the multiple vats located circumferentially on a vertical rotating vat carousel to cross-sectional areas of approximately 4.5-inches by 4.5-inches.

In the MMSL retrofit, spatial constraints limited the multiple vats located circumferentially on a vertical rotating vat carousel to cross-sectional areas of approximately 4.5-inches by 4.5-inches. The limited build size of the retrofitted 250/50 motivated the full development of a new system with multiple material build capabilities comparable to the build envelope of the original 250/50 machine. The new MMSL machine required fabrication of a large system frame, incorporating various 250/50 components and software, and adding a variety of new components and software. By using many existing components and software, the previous engineering development of 3D Systems could be directly applied to this new technology.

By using many existing components and software, the previous engineering development of 3D Systems could be directly applied to this new technology. Components that were transferred from an existing 250/50 to the MMSL machine included the complete optical system (including the optics plate with laser, mirrors, beam expander, scanning mirrors, and focusing lens), the rim assembly (including the laser beam profilers), the associated controllers (computer system, scanning mirror controller, power supply-vat controller) and the wiring harness. In addition to the new frame, the MMSL machine required the development of a new rotating vat carousel system, platform assembly, multi-pump filling/leveling system, and a custom LabVIEW® control system to provide automated control over the MMSL process.

The overall operation of the MMSL system was managed using the LabVIEW® program, which also included controlling a new vat leveling system and new linear and rotational stages, while the 3D Systems software (Buildstation 4.0) was retained for controlling the laser scanning process. As a demonstration of MMSL technology, simple multi material parts were fabricated with vertically and horizontally oriented interfaces. The fully functional MMSL system offers enormous potential for fabricating a wide variety of multiple material functional devices.Keywords: rapid prototyping; stereolithography; functional devices; functional integrated layered manufacturing; multiple material stereolithography;1. IntroductionConventional rapid prototyping (RP) or layered manufacturing (LM) technologies are flexible in manufacturing complex three-dimensional (3D) freeform objects, although these technologies typically have access to only a single build material during fabrication.

As RP technologies have evolved, the materials available for use in RP have advanced considerably where these new materials are enabling LM to be used in rapid manufacturing of end-use products instead of simply producing prototypes. The first commercialized RP technology,624stereolithography (SL), continues to be one of the most widely used LM processes for manufacturing prototypes. In SL, an ultraviolet laser beam is used to scan the surface of a photo- curable polymer that builds 3D parts by selectively solidifying thin layers of the polymer and stacking these layers together. The number of resins available for use in SL has steadily been increasing with several functional high temperature and high strength resin options today.

The number of resins available for use in SL has steadily been increasing with several functional high temperature and high strength resin options today. As a result of these new materials, we believe a market exists for using multiple materials in SL during the same build (including, for example, embedding different colors in medical models or depositing conductive media within mechanical structures to manufacture unique 3D electronic devices). Our particular motivation for this new technology is in the area of tissue engineering where we are using multiple polymers to tailor the bioactive and mechanical properties for particular tissue engineering applications (see Arcaute et al., 2005a, 2005b, 2006 as examples).Our group embarked on developing a multiple material SL machine by first retrofitting an existing 3D Systems 250/50 SL machine with multiple vats positioned on a rotating vat carousel (Wicker et al., 2004).

To build out of multiple materials, the process proceeds as follows. First, the platform is submerged in one of the build vats where building proceeds normally. Different materials are used by simply raising the platform out of the current vat and rotating a different vat underneath the platform providing access to a second material. The process also includes a cleaning step where the platform is submerged and cleaned in a cleaning vat prior to being submerged in subsequent material vats. The process can proceed indefinitely to build a wide variety of multiple material parts.MMSL technology is one aspect of a concept we refer to as Functional Integrated Layered Manufacturing (FILM), where access to virtually any subtractive or additive manufacturing technology, materials, and software is provided in an automated building environment for manufacturing unique functional devices.

In addition to MMSL, our group has integrated a direct write (DW) fluid dispensing system with SL for automated dispensing of conductive media combined with SL-fabricated structures (see Palmer et al. 2004; Medina et al., 2005; Lopes et al., 2006 as examples). This automated machine demonstrates the ability of SL to be combined with non-SL technologies (i.e., using/combining non-UV curable media) for fabricating functional devices, where we have used this machine to fabricate 3D, high density circuitry with integrated embedded electronics. The concept of combining different materials in the same build is not new and many researchers are exploring ways to exploit the advantages of additive manufacturing in producing multi-material parts. For example, Valerio et al.

For example, Valerio et al. (2005) demonstrated a multi material and multi process deposition method to produce parts with specific requirements that could not be met by a single material, such as with graded composition and locally controlled properties. Kumar et al. (2004) presented a concept for multi-material solid freeform fabrication of heterogeneous components using selective laser sintering. Jafari et al. (2000) demonstrated fused deposition of multiple ceramics for fabrication of advanced functional ceramic and composite components. In the end, many of these technologies may be combined in the integrated FILM environment to fully capitalize on individual strengths afforded by each LM technology for producing useful products.We previously demonstrated multi-material fabrication of PEG hydrogel multi-lumen nerve regeneration conduits and functional devices with embedded electronics applying the MMSL concept.

However, in the MMSL retrofit, spatial constraints limited the multiple vats located circumferentially on a vertical rotating vat carousel to cross-sectional areas of approximately 4.5- inches by 4.5-inches. The limited build size of the retrofit motivated the development of a new625stand-alone system with multiple material build capabilities comparable to the build envelope of the original 250/50 machine. This new machine, described here, required fabrication of a large frame to house the system, incorporating various 250/50 components and software, and adding new components and software as required for the MMSL machine to function. The 3D Systems components integrated in the new system were all associated with the beam scanning system and its operation. Additional components and software were required for MMSL operation, where an overall process control software management system was developed using LabVIEW.

Additional components and software were required for MMSL operation, where an overall process control software management system was developed using LabVIEW. The following sections describe the MMSL hardware and software systems followed by demonstrations of MMSL operation where simple multiple materials parts with vertically and horizontally oriented interfaces were fabricated.2. Multiple material stereolithography machine designThe new MMSL machine, as shown in Figure 1, consists of a control center and manufacturing center. The manufacturing center includes the scanning mirror system, laser, and the rim assembly retained from the 250/50 system and a newly developed rotating platform, rotating vat carousel consisting of four vats, vat cover, and rotary stages. The control center includes a multi-pump filling/leveling system, 3D Systems software, overall process control software management system, scanning mirror controller, and signal conditioning module.

The control center includes a multi-pump filling/leveling system, 3D Systems software, overall process control software management system, scanning mirror controller, and signal conditioning module. The new MMSL machine is comprised of hardware that deals with the fundamental component design (including MMSL machine frame, rotating vat carousel system, platform assembly, multi- pump filling/leveling system, and optical system) and software (including 3D Systems software, and a custom LabVIEW® program) to function as a separate stand-alone manufacturing setup.

The vat cross-section (8-inches by 9-inches) was limited by the load carrying capabilities of the rotary stage used for the vat carousel system which in turn constrained the MMSL machine frame to a 4-ft square cross-section. The aluminum extrusions available in 8-ft sections (maximum available length) are utilized to form the MMSL machine frame (4-ft by 4-ft cross-section) while providing the most compact design for the MMSL machine. This square machine setup offers scalability by integrating with other non-SL technologies, as presented by Wicker et al. (2004). A structural analysis of the MMSL machine frame was performed considering the entire load (including laser head and laser power supply, optical system, rim assembly, rotary and linear stages, vat carousel assembly including vats filled with resin, and platform assembly), yielding a maximum deflection of 0.006 inches.

In order to accommodate this deflection and ensure the rigidity of the frame, tee connectors and 45º support brackets were used in each corner as additional support members. The frame was enclosed with transparent, ultra violet (UV) protective Lexan® doors (GE Structured Products Department, Mt. Vernon, Indiana) on three sides and another UV-protective Lexan® sheet on the back for easy access to all the components of the new MMSL machine and to protect from the UV radiation of the laser. A 6061 aluminum plate (0.5-inches thick) was used as a base of the manufacturing center to accommodate the rotating vat carousel system. The optical system retained from the original 250/50 machine was placed on another aluminum plate (0.25-inches thick) located on top of the MMSL machine.Resin vatVat carousel frameShaft for carouselShaft holding plateRotary stage(a) (b)Figure 2.

(a) Vat carousel assembly exploded view, (b) Vat carousel assembly.2.2 Rotating vat carousel systemThe rotating vat carousel system includes four stainless steel vats (8-inches by 9-inches), located circumferentially, as shown in Figure 2. The maximum build envelope of the MMSL machine based on the vat design was 6.5-inches by 6.5-inches by 4.75-inches. The vertical build was limited to 4.75-inches due to the platform supporting plates, the slope provided to the resin vats, the positioning and the travel (500 mm) of the linear stage (Z-stage). The maximum vat cross-section of the MMSL machine was limited, as compared to the original 250/50 system vats (approximately 11.5-inches by 14.5-inches), by the torque output of the Aerotech rotary stage (425N-m, which is the highest capacity within the Aerotech range of stages). The volume of the627MMSL resin vats was 9 liters as compared to the 32.21 liters volume of the original 250/50 machine vats.

The volume of the627MMSL resin vats was 9 liters as compared to the 32.21 liters volume of the original 250/50 machine vats. The vats were rotated about a vertical axis (the axis of a 2-inch diameter aluminum shaft) in order to position a specific vat below the build platform. ADRT-200 direct drive rotary stage (Aerotech Inc., Pittsburg, Pennsylvania) was selected for the vat carousel system due to its high accuracy (±30 arc-sec) and high torque output (with in the Aerotech range of stages), along with superior angular positioning, velocity control, and load capacities (30-173 kg). Finite element analysis, using COSMOSWorks®, available from SolidWorks® Education Edition 2005-2006, was performed on each vat to calculate the vat deflection under static loading conditions. Considering maximum volumetric capacity condition, fully restrained flange bottom faces, and hydrostatic pressure, the analysis resulted in a maximum deflection of 0.0014-inches for each vat.

Considering maximum volumetric capacity condition, fully restrained flange bottom faces, and hydrostatic pressure, the analysis resulted in a maximum deflection of 0.0014-inches for each vat. Stainless steel (0.0625-inches) vats were fabricated to accommodate the resulting deflection and to store different resins for multi-material fabrication. A vat cover (ABS plastic) fabricated using the FDM machine was designed to avoid external resin contamination and also protect the resin in other vats from laser scattering and adverse environmental effects.

The vat carousel frame was composed of a 6061 aluminum plate at the center and a series of aluminum support columns (1 square inch cross-section) attached to it, in order to accommodate the multiple vats filled with resin.2.3 Platform assemblyA stainless steel build platform (6.5-inches by 6.5-inches in cross-section), as shown in Figure 3, was designed to fit into the vats while maintaining the same spacing between the platform and the sides of the vat as in the original 250/50 system. The build platform also provides space within the vat for the level sensing floating device. The platform assembly was mounted on the ADRS-200 direct drive rotary stage (Aerotech Inc., Pittsburg, Pennsylvania) in order to provide for angled part building, as well as washing and drying of multiple material parts.

The rotary stage was further secured to a high precision Z-stage (Model # ATS10050, Aerotech Inc., Pittsburg, Pennsylvania) with high accuracy (±20μm) and repeatability (±1μm), as compared to the original 250/50 machine (accuracy of ±50μm and repeatability of ±7.6μm) while controlling the build layer thickness of the part. Thus, the platform can be manipulated vertically as well as at any given angle about a horizontal axis for customized part fabrication.Rotary stageRotating platformExtension postsMechanical brake(a) (b)Figure 3. (a) Platform assembly exploded view, (b) Platform assembly.To maintain the position of the platform during part building, a pair of mechanical brakes was designed and attached to the rotary stage as shown in Figure 3b.

A custom-made mechanical brake628consisting of a pair of small aluminum plates was manufactured to accommodate the shape of the circular disc of the rotary stage placed in a milled slot in the rectangular aluminum plates (5-inches by 2.5-inches in cross-section). The rotary stage was locked in position when the brake was activated. A pair of 9.5-inches aluminum extension posts (1-inch diameter) was attached to the rotary stage via a rectangular plate to allow the laser incidence center to be coincident to the center of the platform and to position the platform in the scanning range of the optical system. In addition, recoating technologies can also be employed in this machine by incorporating them individually in each vat or a conventional recoating system for multiple vats.

In addition, recoating technologies can also be employed in this machine by incorporating them individually in each vat or a conventional recoating system for multiple vats. Although the stratified part fabrication and the space constraints due to integration of a new rotating vat carousel assembly eliminated the use of a ZephyrTM blade (sweeping during part build), a dip coating approach with 0.008-inches layer thickness was used to demonstrate MMSL fabrication. The ZephyrTM blade in the 250/50 system allows for resin surface uniform coating with 0.004-inches thick layers as compared to resin leveling by surface wetting alone. The elimination of the ZephyrTM blade required the use of dip coating approach that limited the vertical resolution of the MMSL system.

The elimination of the ZephyrTM blade required the use of dip coating approach that limited the vertical resolution of the MMSL system. However, research is ongoing to incorporate a new recoating system that will allow building with higher resolution.2.4 Multi-pump filling/leveling systemThe auto leveling system of the existing 250/50 machine was replaced by an automatic multi- pump filling/leveling system (including laser displacement sensor, peristaltic pump, pump heads, and flexible tubing) to regulate the resin level in the vat. The multi pump filling/leveling system provided high accuracy (±5μm) control for maintaining resin levels in the multiple vats of the MMSL machine. A Masterflex® L/S peristaltic pump (±13μm resolution), as depicted in Figure 4, facilitated adding and removing desired quantities of resin to and from each vat and allowed isolation of the resin from the moving parts of the pump.

The most significant advantage of the peristaltic pump is the use of flexible tubing (MasterFlex® Tygon® L/S 18) for the multi-pump system, as the resin being pumped remains inside the tubing at all times. This feature greatly reduced the risk of contamination by offering complete control over the content and purity of the resin along with the integrity of the process. This resulted in minimizing the maintenance, cleanup, and resin changeover times while enabling the pump to deliver various fluids simply by changing the tubing within the pump heads. Four pump heads (Model # 77201-60, L/S Easy- Load® II, Cole-Parmer Instrument Company) were used to facilitate the fabrication of the functional devices with multiple materials.

Four pump heads (Model # 77201-60, L/S Easy- Load® II, Cole-Parmer Instrument Company) were used to facilitate the fabrication of the functional devices with multiple materials. A sensing head (including a solid state laser and a position sensitive detector) mounted on the rim assembly is utilized to determine the resin surface level using a non-contact laser triangulation measurement system.Peristaltic pumpPump headLaser displacement sensorFlexible tubingFigure 4. Multi pump filling/leveling system.629One of the issues encountered by Wicker et al. (2004) with the retrofit design was the re- registration of the part to maintain a uniform level of resin on the build layer in multiple vats to produce multi-material parts. This was addressed using a wet film coating thickness gauge (Elcometer®) to manually check the resin level on the build layer of the part as demonstrated by Sandoval et al. (2005).

This was addressed using a wet film coating thickness gauge (Elcometer®) to manually check the resin level on the build layer of the part as demonstrated by Sandoval et al. (2005). The multi pump filling/leveling system of the new MMSL machine addressed this issue by employing a laser displacement sensor (MicrotrakTM 7000) and a sensing head coupled to a floating device calibrated for the various fluids’ buoyancy parameters. A rubber cellulose floating device (3/64 inches thickness) was essential due to the difference in viscosities (WatershedTM: 260 cps at 30ºC, and NanoformTM: 570 cps at 30ºC) (DSM Somos®, New Castle, Delaware) and the appearance (WatershedTM: Optically clear, and NanoformTM: Opaque gray) of the resins used in SL.

The calibrated values (WatershedTM: 0.012-inches and NanoformTM: 0.018-inches above resin level) were accommodated in the LabVIEW® program for automatic leveling of various resins, by calculating the error and adjusting the resin level surface to a calibrated (zero) value. A PID (Proportional Integral Derivative) control was used for automatic resin leveling, due to its functional simplicity and robust performance in a broad range of operating conditions. A preliminary experiment using varying PID readings aided in determining the values for WatershedTM and NanoformTM resins to automatically maintain uniform levels in multiple vats.2.5 Transferred technologiesSeveral components of the existing 250/50 system were retained (including the optical system, rim assembly, the associated controllers and the wiring harness) in the new MMSL machine.

The components required for controlling the laser scanning process in the original 3D Systems machine and any components required to allow the scanning system to function were retained. The optical system (including the optics plate with laser, mirrors, beam expander, scanning mirrors, and focusing lens) was transferred to utilize the advantages of the existing technology and thus maintain the original SL operation functional. The rim assembly (including the laser beam profilers), and the associated controllers (computer system, scanning mirror controller, power supply-vat controller) were used to achieve the functionality of the laser scanning process. The MMSL machine used a 355 nm solid state laser (Lightwave® Electronics, Mountain View, California) due to its high resolution and small beam diameter (0.004-inches).

The MMSL machine used a 355 nm solid state laser (Lightwave® Electronics, Mountain View, California) due to its high resolution and small beam diameter (0.004-inches). The wiring harness was utilized to initialize the 3D Systems software and achieve successful interface with the custom LabVIEW® control system by acquiring the analog signal from the stepper motor of the recoating system of the 250/50 machine. The software (including 3D Systems software, and a custom LabVIEW® program) required to provide automated control over the hardware is described in the following section.3. Multiple material machine operation overviewThis section highlights the software system of the new MMSL machine. The development of an overall process management and control software to control the MMSL operation is the most significant element of this research. A generalized multiple material fabrication process is depicted in Figure 5.

A generalized multiple material fabrication process is depicted in Figure 5. The process is initialized by submerging the build platform in the vat containing the first resin material (WatershedTM). Normal SL build is performed to fabricate the desired part with first resin. When finished, the platform is raised, and a cleaning vat is rotated underneath the platform so that the platform and part can be cleaned. Once cleaned and dried, the platform is raised out of the cleaning vat and the second material vat is rotated underneath the platform (NanoformTM). The SL build continues with the second material until completion, and630the process continues indefinitely to fabricate fully functional multiple material parts. The entire MMSL process is managed by the overall process management and control software using a LabVIEW® program.

The entire MMSL process is managed by the overall process management and control software using a LabVIEW® program. The software controls the rotating vat carousel system, new linear and rotational stages, the filling/leveling system, and communicates directly with the 3D Systems software (Buildstation 4.0) for controlling the laser scanning process.Build using material 1 (raise platform when completed)Rotate cleaning vat under build platformRotate material vat 1 under build platformFinished multiple material partClean and dry (raise platform when completed)Continue the process to build multiple material functional devicesBuild using material 2 (raise platform when completed)Rotate material vat 2 under build platformFigure 5. Multiple material stereolithography (MMSL) operation.The key elements for successful operation include being able to communicate between these two systems.

Multiple material stereolithography (MMSL) operation.The key elements for successful operation include being able to communicate between these two systems. A signal conditioning module (SC-2345 carrier with configurable connectors, National Instruments, Austin, Texas) is used to acquire the analog signal (change in voltage) from the stepper motor of the recoating system used in the 250/50 machine after laser scanning of each part layer. The custom LabVIEW® program uses this change in voltage and opens an interlock switch (a part of the wiring harness retained from the 3D Systems machine) thereby momentarily stopping the 3D Systems program to perform the dip coating operation.

The sequence continues as shown in Figure 6 for successful MMSL fabrication.631Move building platform to home positionMove vat carousel to home positionMove platform to the start positionSensor positionResin levelingInterlock switch closeVoltage measurementAnalog signal from the stepper motorSurface wetting time (Z-wait)Any change in voltageNoYesResin levelingDelayDip coating of the build platformInterlock switch openSensor positionPause processOpenClosedInterlock switchPre-fabrication positioningCustom LabVIEW® programAnalyzingDrawingRecoating3D Systems programFigure 6. Process flowchart of custom LabVIEW® program.Any CAD package (such as SolidWorks®) can be used to design the individual parts and export them into STL file format.

Process flowchart of custom LabVIEW® program.Any CAD package (such as SolidWorks®) can be used to design the individual parts and export them into STL file format. 3D Systems proprietary slicing software (3D Lightyear™ 1.4) was used to convert the STL files generated during the part preparation and design stages into the build file format required for the operation of the MMSL machine. A process management and control software (described above) was developed using LabVIEW® and its process flowchart is depicted in Figure 6. MMSL operation is divided into three distinct sections (including pre- fabrication positioning, LabVIEW® program, and 3D Systems program). The pre-fabrication positioning starts with initializing the new platform assembly (secured to the Z-stage) and the632rotating vat carousel assembly (representing the X-stage) to attain respective home positions.

The platform is lowered to the start position and a Masterflex® L/S peristaltic pump is activated to level the resin surface with the platform. After the pre-fabrication positioning is complete, the 3D Systems program is commenced to scan the first layer of the desired part. The LabVIEW® program opens an interlock switch, after the laser beam scans a layer, to pause the 3D Systems program and commands the build platform to traverse vertically downward by a distance equal to the build layer thickness using a dip coating approach. The multi-pump filling/leveling system is then activated to maintain 0.008-inches of the resin on the build layer of the desired part. The LabVIEW® program then closes the interlock switch to commence the laser scanning process. The traditional SL part building is performed layer by layer by integrating the LabVIEW® program with the 3D Systems software to produce desired part with the first resin.

The traditional SL part building is performed layer by layer by integrating the LabVIEW® program with the 3D Systems software to produce desired part with the first resin. Once the part using the first resin is completed, the platform is raised out of the vat and the LabVIEW® program is used to command the vat carousel to rotate to the cleaning and drying vat for cleaning. After desired cleaning of the part and the platform is completed, it is immersed in the next resin vat to continue MMSL fabrication. The process is repeated indefinitely to produce multiple material functional devices. The LabVIEW® program thus controls the interlock switch that acts as an interface between the new MMSL system and the original 250/50 system facilitating successful MMSL operation. The following section demonstrates the manufacturing capabilities of the new MMSL system as compared to the existing 250/50 system.4.

The following section demonstrates the manufacturing capabilities of the new MMSL system as compared to the existing 250/50 system.4. Fabrication demonstrationSimple parts with a horizontally oriented interface (a vertical stratified build) and vertically oriented interfaces (a horizontal stratified build) were manufactured using DSM Somos® WatershedTM and NanoformTM resins due to their considerably different SL resin properties. The part with the horizontally oriented interface is shown in Figure 7 while the part with the vertically oriented interfaces is shown in Figure 8. WatershedTM is the most popular resin and was selected due to its low viscosity while producing strong, tough, water-resistant parts. NanoformTM (a nanoparticle filled material) was selected due to its ability to withstand high temperatures and fabricate strong, rigid parts.

NanoformTM (a nanoparticle filled material) was selected due to its ability to withstand high temperatures and fabricate strong, rigid parts. Studies can be conducted to investigate the interfacial bonding between these two resins and various other resins could be included for future research. As shown in Figures 7 and 8, the MMSL machine successfully fabricates multiple material parts with horizontal or vertical material stratification. Some of the important aspects to be studied are interfacial bonding between different SL materials which could expand the scope of MMSL technology into various engineering fields.NanoformTMHorizontally oriented interfaceWatershedTMSupports(a)(b)Figure 7. (a) Vector file demonstrating vertical stratified build, (b) Finished vertical stratified part633WatershedTMVertically oriented interfaceNanoformTM(a) (b) Figure 8. (a) Vector file demonstrating horizontal stratified build, (b) Finished horizontal stratified part5.

(a) Vector file demonstrating horizontal stratified build, (b) Finished horizontal stratified part5. ConclusionsA novel MMSL setup was developed to fabricate multiple material functional devices. The new MMSL setup operates similarly to the existing 250/50 system, although a custom LabVIEW® management and control system was used to manage the overall process. The 3D Systems software was retained for controlling the laser scanning process. The new MMSL machine was composed of two basic systems: a hardware system (including machine frame, rotating vat carousel assembly, platform assembly, multi-pump filling/leveling system, and scanning system) and a software system (including 3D Systems software and custom LabVIEW® program).

Several components were retained from the existing 250/50 system required for laser scanning and new components and software were developed to add MMSL functionality.The new MMSL setup was successfully used to manufacture simple multiple material parts featuring vertically as well as horizontally oriented interfaces. This MMSL system offers enormous potential for fabricating a wide variety of multiple material functional structures and research is ongoing to both improve the functionality of the MMSL system and to use the system for exploiting these new MMSL capabilities. We believe a market exists for using multiple materials in SL and our group is exploring a variety of possible opportunities, including multiple colored SL models, embedded electronics and 3D circuitry, and tissue engineering.AcknowledgmentsThe research presented here was performed at UTEP in the W.M. Keck Border Biomedical Manufacturing and Engineering Laboratory (W.M.

Keck Border Biomedical Manufacturing and Engineering Laboratory (W.M. Keck BBMEL) using equipment purchased through Grant #11804 from the W.M. Keck Foundation. This material is based in part upon work supported by the Texas Advanced Research (Advanced Technology/Technology Development and Transfer) Program under Grant Number 003661-0020-2003. Support was also provided through the Mr. and Mrs. MacIntosh Murchison Chair I in Engineering and through research contract 54004 from Sandia National Laboratories in the Laboratory Directed Research and Development (LDRD) program. Sandia National Laboratories is a multi-program laboratory operated by Sandia Corporation, a Lockheed Martin Company, for the United States Department of Energy’s National Nuclear Security Administration under contract DE-AC04-94AL85000.634References 1. 3D Systems®, Stereolithography Buildstation User Guide, 1995, 3D Systems®, Valencia,California.2.

velopment of an economic decision support for the application of Additive Manufacturing in aerospaceG. Deppe*, C. Lindemann*, R. Koch*Direct Manufacturing Research Center (DMRC) and Chair of Computer Application and Integration in Design and Planning (C.I.K.), University of Paderborn, Mersinweg 3, 33098 Paderborn, GermanyAbstractAdditive Manufacturing offers a high potential in aerospace industry due to its freedom of design and the ability to manufacture complex and lightweight parts. The low number of units, high quality standards and fast response time are special challenges that have to be met especially in the Maintenance, Repair and Overhaul sector. Thus, companies have to decide at which point it is economic to apply Additive Manufacturing. However, companies lack experience on this new technology. This is why a tool is required that takes into account the above mentioned crucial points and supports the decision process.

However, companies lack experience on this new technology. This is why a tool is required that takes into account the above mentioned crucial points and supports the decision process. The paper analyzes aviation’s characteristics with regard to Additive Manufacturing. The structure of current MRO repair workflows is investigated to identify a feasible application for Additive Manufacturing. Additionally the supply chain will be examined to indicate the benefit which the technology can generate in this highly demanding field. The findings are integrated into a methodology that supports the decision whether to apply Additive Manufacturing on the basis of costs, time and quality.IntroductionIncreasing product complexity and shorter life cycles are major challenge for industry today. To stay competitive companies have to apply innovative technologies to produce individual and customized products. They have to improve quality, costs and time [Mei14].

To stay competitive companies have to apply innovative technologies to produce individual and customized products. They have to improve quality, costs and time [Mei14]. Due to the layer-based production approach of Additive Manufacturing (AM) this technology offers the opportunity to enhance those key factors. The high flexibility of this technology enables an economic small scale production of highly complex products [Gebh13]. It has the potential to shift the production from a product specific production line to a system that integrates many different production steps. An environment exhibiting high requirements on time, quality and costs can be found within the Maintenance, Repair and Overhaul (MRO) of aerospace parts. The production and repair of spare parts for aircrafts is demanding and conventional technologies are limited in their ability to meet these requirements without using warehouses as a buffer while AM offers a new scope.

[WHY04] Despite AM’s benefits it has to be assessed whether an application is useful for the intended application. Most companies are neither experienced in applying this technology nor in evaluating its utilization. They lack empirical value to calculate the economic impacts in contrast to conventional technologies where a detailed evaluation can be conducted. To compare both, cost drivers for AM have to be identified and associated with further influence factors such as time and quality which are especially important to the aerospace industry [SGF+08]. Therefore an approach is required that takes into account the key factors and determines the most cost-efficient manufacturing strategy. The aim is to recommend an action for an individual decision process. This paper describes a concept for the decision support for metalAM parts on the basis of the special characteristics of aerospace on a single part comparison.

This paper describes a concept for the decision support for metalAM parts on the basis of the special characteristics of aerospace on a single part comparison. The use case is a defect part that is sent to the workshop in order to determine whether to ‘make or buy’ the part and to specify which technology should be used to do so.Aerospace and Additive ManufacturingThe aerospace sector will gain importance as forecasts predict. The aerial sector will show a continuous growth in the next decades that exceeds the global growth rate [Boei12]. Nevertheless the piece number of aerospace parts will remain small. Due to their application field aerospace parts have to meet special requirements. The complex aircraft design often leads to complex parts. Additionally aircrafts have to face extreme environmental conditions that put the part’s requirements on a high level.

The complex aircraft design often leads to complex parts. Additionally aircrafts have to face extreme environmental conditions that put the part’s requirements on a high level. The rising oil price forces the aircraft manufacturers to develop aircrafts with a higher performance and at the same time a lower fuel consumption to reduce the emissions and overall the operation costs. A decrease in the consumption of raw material during production and in the end a lower aircraft weight is of particular importance for the light weight construction. [Bul09] For those reasons the use of special materials and manufacturing technologies is often required. In combination with the high quality standards in aeronautics, high unit costs arise. AM is capable to meet these requirements as it offers a flexible production and can manufacture complex parts at nearly no extra costs. Thus it is possible to create lightweight parts with a low input of raw material leading to a low buy-to-fly ratio.

Thus it is possible to create lightweight parts with a low input of raw material leading to a low buy-to-fly ratio. This is one reason why the aerospace industry is a pioneer for the application of AM. [GEW13] However, as the experience with AM is limited in aerospace, the number of additively manufactured parts in aircrafts is comparably low. The certification of those parts is a difficult and long process as it takes time to proof their durability and safety. Nevertheless, companies and research projects such as “RepAIR” are developing concepts to do so. [Rep-ol] Currently, the focus is on two AM technologies, both metal based: Selective Laser Melting (SLM) by SLM Solutions and Electron Beam Melting (EBM) by Arcam. They are both powder bed based technologies but differ in their energy source. The SLM process uses a laser which is deflected by a scanning unit to melt the metal powder in an inert gas environment to prevent oxidation.

The SLM process uses a laser which is deflected by a scanning unit to melt the metal powder in an inert gas environment to prevent oxidation. For EBM, the energy source for melting the powder is not a laser but induced by an electron beam. The build chamber therefore has to be a vacuum, consequently preventing oxidation as well. The beam is focused and deflected by electric inductors. [Gebh13]MRO Processes and Supply ChainThe high investment in an aircraft implies that the airline tries to maximize the life time of this investment. MRO providers ensure the airworthy condition of aircrafts and their availability by maintaining all of its components which cause 10-20% of the airline’s overall costs for operating an aircraft [Mens13a]. If a defect is detected then the component is sent to the workshop in order to be repaired if this is feasible.

If a defect is detected then the component is sent to the workshop in order to be repaired if this is feasible. The repair process usually exhibits a generic structure [Mens11] of:a) inspection b) cleaning c) preparing for applying material d) applying material e) finishing the part’s geometry f) coating g) final testingIn this context AM can be used instead of build-up welding, enabling a much more accurate geometry re-establishment which decreases the effort of further subsequent steps. This is why the whole process has to be taken into account to assess the suitability of an AM application. A detailed plan to include AM cannot be given, the assessment whether AM is able to compete in time and costs with current repair processes has to be conducted on a single case basis. Besides the economic analysis whether AM can compete in terms of costs with welding for the production or repair, it is also important to consider possible alterations in the supply chain.

Besides the economic analysis whether AM can compete in terms of costs with welding for the production or repair, it is also important to consider possible alterations in the supply chain. The aerospace industry has high demands on the spare part supply chain. It has to be flexible, fast and efficient at the same time to be able to deliver the required part in time to the point of use. In combination with the high value of aerospace parts and the long product life cycle this leads to complex and expensive logistic concepts as part availability and part quality are the major requirements within spare part logistics. [SGF+08] [WHY04] By nature, the aerospace industry is a global market and there are often labor intensive maintenance checks which are more likely to be conducted in low wage countries which enforce intense competition. Thus, European MRO providers have to search for application fields where they can offer their services competitively.

Thus, European MRO providers have to search for application fields where they can offer their services competitively. With the help of AM the supply chain process can become more efficient [Fish97]. The batch production and the high set up time cause that conventional manufacturing technologies are not flexible enough to react to a short-term demand. To do so, intensive warehousing is necessary which generates high expenses only due to capital lock-up besides further cost drivers [FrLi09]. Especially parts that are considered as slow-moving can benefit from AM. A conventional batch production is not economic so that a single part can be produced by AM when it is required instead of manufacturing an entire batch and store the surplus parts for a long time until they are inquired.Production Cost CalculationThe cost calculation is characterized by specific influence factors. It is important to identify those so that they can be evaluated and controlled.

It is important to identify those so that they can be evaluated and controlled. The cost drivers can be separated into two groups as figure 1 shows: fixed costs and variable costs. The acquisition of the AM machine and investment in further tools and machines are fixed costs. The amount and extent highly depends on the available facilities and the chosen machine configuration. Costs for maintenance and overhaul have to be taken into account additionally. Occupancy- and labor costs are also independent of single build jobs. One major cost driver of the variable costs is the material usage which depends on the part design. Further variable costs are the operating expenses which are caused by energy and inert gas for the production to prevent the material’s reaction with oxygen and a proportion of the overhaul costs.

Further variable costs are the operating expenses which are caused by energy and inert gas for the production to prevent the material’s reaction with oxygen and a proportion of the overhaul costs. [Gebh13]Figure 1: Relevant elements for the cost calculation [Gebh13]The main influence factors are the build-time, batch size and build chamber utilization rate, the workload, material consumption, process stability, technical progress and service (see figure 2). The build time is one of the key cost drivers and an essential criterion for the assessment of the suitability of this technology for a certain part. It incorporates secondary time for the pre-processing and the set-up of the build job and the AM machine. Due to this fact a high degree of build chamber utilization is preferable to minimize the unit costs as the standard pre- and post-processing activities such as cleaning the machine can be apportioned to more than one part.

Nevertheless AM is capable of an economical single part production as the set-up time is considerably lower than the one for conventional manufacturing. In general, the production costs do not decrease proportional to the build chamber utilization rate. Support structures also have to be taken into account, they depend on the part geometry and the chosen material as well as the orientation of the part in the build chamber and increase the material costs. Especially the latter one is another key factor for the production costs as it influences significantly the build and post-processing time. [Gebh13] [LJM+12]Figure 2: Influence factors of Additive Manufacturing [Fre15]The flexibility of AM offers the opportunity to use it for several different applications. To remain the flexibility the machines cannot run at full capacity, however, short-term production orders cannot guarantee an optimal workload which is required to be cost- efficient.

To remain the flexibility the machines cannot run at full capacity, however, short-term production orders cannot guarantee an optimal workload which is required to be cost- efficient. A trade-off strategy has to be chosen in order to get to an optimal approach. [KPH13] While AM material is currently comparatively expensive the technology is characterized by a high utilization rate which often compensates the material costs. Additionally, metal powder that has not been solidified during a production run can be recycled and used again [Gebh13]. A critical topic is still the process stability. Quality problems can lead to multiple iterations of a build job and thus leading to high costs. Currently, high efforts for assuring the quality have to be undertaken. Future developments in this field will lead to a much higher part reliability [CCB+14]. This is why the technical progress in the field of AM and its machines is another aspect that has to be considered [LJM+13].

This is why the technical progress in the field of AM and its machines is another aspect that has to be considered [LJM+13]. To be able to produce with the most recent technology, high investments are necessary as the technology is fast evolving and updates of hardware and software are usually necessary from time to time. The flexibility of the technology also depends on the service level of the machine manufacturers. If a repair is necessary and neither skilled staff nor spare parts are available, the downtime can lead to unexpected costs and influences the utilization rate negatively.Decision AlternativesIn order to develop a decision support, the decisions that can be made have to be known. Therefore, the general, operational procedures of MRO provider have to be taken as a basis. If a defect part is inspected at the workshop and classified as not repairable, a new one has to be ordered from the OEM.

If a defect part is inspected at the workshop and classified as not repairable, a new one has to be ordered from the OEM. If a repair is possible it has to be assessed which resources in terms of material, tools and personnel are required. If an economic repair solution can be found, a work order is started. If not, it is checked whether the repair can be outsourced or stored to repair it at a later point of time. [Mens13b] The decision support in this paper focuses on the selection between three alternatives:- production of a conventional milling part - the acquisition of a new partin-house repair applying AMThe aim is to develop a concept to calculate a cost-efficient repair. The concept supports the decision process as a standardized instrument to assess the costs of the decisions alternatives and overall saves time and costs during this process.

The concept supports the decision process as a standardized instrument to assess the costs of the decisions alternatives and overall saves time and costs during this process. On a strategic level it has to be evaluated if the investment into a new technology is useful to establish a new repair solution.Configuration relevant subsystemsTo calculate the costs for the AM production it is necessary to provide information for different aspects. This is why different subsystems have to be defined. They are divided into four configuration units as figure 3 shows.Figure 3: Overview of the configuration units for the Decision ComponentThey represent separate configuration units which have to be filled by the operator in order to allow the calculation. This process can be supported by pre-defined master data that can be overruled but enables an easier operation of the tool.

This process can be supported by pre-defined master data that can be overruled but enables an easier operation of the tool. The system has to consolidate the given inputs and to ensure that all calculation relevant data is available.Machine configuration All machine specifications have to be entered in this section. As milling and AM machines differ in their characteristics this has to be done for each separately. The dimensions of the build chamber have to be known so that it can be determined which parts can be repaired. For milling machines the tools have to be indicated. To detail the cost calculation the power consumption can be specified which fosters further potential analyses of the ecological effects.Material configuration The configuration of the material is done accordingly. Material for milling and AM can be different if a certain very specific material is not available for one technology, but can be substituted by a similar one.

Material for milling and AM can be different if a certain very specific material is not available for one technology, but can be substituted by a similar one. The main focus is on the costs that arise for the repair of a part which is often significant for AM due to the high material prices but can be balanced because of the low buy-to-fly ratio. For AM additional information about the density and a waste factor are required for the calculation. For milling, only the block of solid material is incorporated, using the bounding box as a reference. Information regarding the supply chain management can be stated, additionally. The availability of raw material has to be ensured. If it is not in stock it has to be ordered from the OEM requiring a certain amount of time which has to be considered when calculating the most cost-efficient repair considering also the time frame.

If it is not in stock it has to be ordered from the OEM requiring a certain amount of time which has to be considered when calculating the most cost-efficient repair considering also the time frame. As especially in aerospace time often equals money, this can significantly influence the decision.Part configuration The part configuration is not divided into the two categories for AM and milling. Parts can be chosen from a database or they can be described manually by key data required for the calculation. That includes the dimensions of the part in order to assess whether the part fits into the build chamber and the quantity of other parts that can be produced along with it. This is important for the assessment of the machine capacity. The operator has to choose the x,y,z direction accordingly to the part’s orientation within the build chamber. The height that has to be reestablishing and the correspondent build up part volume are the major production time determining variables.

The height that has to be reestablishing and the correspondent build up part volume are the major production time determining variables. Thus, they are one of the major cost drivers of the AM calculation. As the prediction of the build-time on the basis of height and part volume is still not accurate for all machines and materials, this can also be filled in manually if a more accurate value is available from former production runs. Additional programming costs can be stated as well. Besides from that a complexity factor has to be chosen for pre-processing the AM build job and for the post-processing effort. There are different ones to choose from. It is differentiated between new and known build jobs and additionally between simple and complex parts all influencing the effort that has to be taken into account for the processing time.

It is differentiated between new and known build jobs and additionally between simple and complex parts all influencing the effort that has to be taken into account for the processing time. Simple and known build jobs can be prepared within a few minutes while complex parts often require a much higher effort for determining orientation and necessary support structure. In terms of the supply chain management the costs for the acquisition of a new part from the OEM can be specified as well as the expected delivery time to compare not only the costs but also the time for the repair and the new procurement.Others/ general configuration This subsystem encompasses all other required data which depend on the company. Again, AM and milling are considered differently allowing the separate definition of costs for the two technologies. They can be divided inI. II. III. maintenance and production overhead costs IV.

Again, AM and milling are considered differently allowing the separate definition of costs for the two technologies. They can be divided inI. II. III. maintenance and production overhead costs IV. V.labor expenses and operating costs hard- and software costslevel of capacity depreciationThe labor expenses can be assigned to the manufacturing steps pre-processing, manufacturing and post-processing and are subdivided into wages for blue collar worker, skilled worker and engineers. It is furthermore defined how much manual work each step requires being theninfluenced by the complexity factor. Investments for the purchase of machines and software as well as their maintenance expenses have to be filled in. In combination with the annual utilization of the machines and the depreciation time frame the hourly rate for the machines is fixed.

In combination with the annual utilization of the machines and the depreciation time frame the hourly rate for the machines is fixed. After those inputs have been made once they only have to be updated from time to time as they are not depending on the part or the machine and remain stable for a longer period of time.The presented system is now able to calculate the costs for an AM production and the milling of a part while taking into account the supply chain aspect whether one option will take longer than the other.Evaluation subsystem To provide information as a basis for the decision of a repair or procurement a detailed, yet significant evaluation has to be available. A graphical overview of the calculation results facilitates the decision showing which cost driver is dominating each technology. The share of each of them with regard to the overall costs is important to identify potentials and deficits in the process for further improvements.

The share of each of them with regard to the overall costs is important to identify potentials and deficits in the process for further improvements. The illustration of the results into quality, costs and time graphs reflects the fundamental approach every economic oriented company requires. Especially for MRO service providers these three aspects are crucial for conducting their business successfully [Mei14]. The graphs enable the examination of the relations between quality, costs and time (QCT) for the different choices taking more than just the costs into account. They are displayed in percent with 100% for the best rated process option of all calculated repair alternatives. The QCT graphs show clearly which repair alternative offers the best overall solution in terms of cost, quality and processing time.Figure 4: Evaluation of the calculation resultsAnother important aspect for aerospace is the documentation of this process to foster the transparency and the confirmability.

Therefore a document has to be generated that contains all input from the operator and the databases and additionally all calculation results that the tool compiles in order to be able to reproduce the decision process. It furthermore supports the improvement of the tool as the output can be matched with data of the actual production process. This is especially relevant for AM as for example the build time estimation has to be enhanced.Summary and outlookDue to the specific characteristics of aerospace, Additive Manufacturing is suited to be applied in this industry. The required flexibility for low quantity and highly complex products cannot be realized by conventional technologies without reverting to extensive warehousing. The usefulness of applying AM for a certain use case yet has to be proofed. Therefore a methodology is required that supports the decision process.

The usefulness of applying AM for a certain use case yet has to be proofed. Therefore a methodology is required that supports the decision process. Especially because companies are not experienced in assessing the production costs of AM and additional benefits have to be taken into account to fully exploit the benefits AM offers. Based on identified key cost drivers four configuration units have been set up. They provide a standardized process to gather the data that is required for the calculation of the repair choices which is supported by predefined data but can always be adjusted manually. The tool then calculates the expected costs for AM and milling and the procurement of a new part from the OEM. The evaluation is illustrated by charts showing the share of each cost driver from the overall costs. Additionally, quality, costs and time graphs provide information of the key elements of MRO business showing the overall best solution for the defect part.

Additionally, quality, costs and time graphs provide information of the key elements of MRO business showing the overall best solution for the defect part. Thus, the concept for the decision component allows the monetary assessment of repair processes including AM. The documentation of the complete input and output data fosters the transparency and traceability of decisions which is a crucial aspect in aerospace. For future work a detailed comparison of sample parts is required in order to assess the validity of the tool. Therefore the tool has to be improved in the evaluation of costs, quality and time to allow a detailed analysis of its functionality. It furthermore can be enhanced by strategic levels, calculating the product life cycle costs and an ecological investigation.

It furthermore can be enhanced by strategic levels, calculating the product life cycle costs and an ecological investigation. The aerospace specific tool could also be adapted to the needs and specifics of other industries in order to proof its general applicability.AcknowledgementsThe research leading to these results has received funding from the European Union Seventh Framework Programme (FP7/2007-2013) under Grant Agreement n°605779 (project RepAIR). The text reflects the authors' views. The European Commission is not liable for any use that may be made of the information contained therein. For further information see http://www.rep-air.eu/.Literature[Boei12]Boeing (2012): Current Market Outlook 2015-2034. Hg. v. Market Analysis. Boeing. Seattle, WA. URL: http://www.boeing.com/commercial/market/long-term-market/[Bul09]Bullinger, Hans-Jörg (2009): Technology Guide: Principles Applications – Trends. Springer, Berlin.[CCB+14]S. Clijsters; T. Craeghs; S. Buls; K.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDevelopment of an Engineering Diagram for Additively Manufactured Austenitic Stainless Steel AlloysZachary T. Hilton, Joseph W. Newkirk, Ronald J. O’MalleyDepartment of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractAustenitic stainless steels are the most widely applied types of stainless steels, due to their good weldability and high corrosion resistance. A number of engineering diagrams exist for the purpose of providing insight into the behavior of these steels. Examples of these diagrams are constitution diagrams (aka Schaeffler Diagrams) which are used to approximate the solidification path of the alloy and the amount of retained ferrite in the solidified matrix.

Other diagrams are the Suutala diagram, which approximates cracking susceptibility, and microstructural maps, which predict the solidification path by varying a processing parameter, such as cooling rate. By combining these diagrams, a much more concrete conclusion can be made as to the behavior of a particular steel. This approach could be used to determine differences in behaviors between two different compositions. The developed diagram would be intended for use with rapid solidification phenomena as observed in the selective laser melting process.IntroductionConstitution diagrams such as the Schaeffler diagram have long been used to predict the microstructural behavior of stainless steels in welding processes [1]. As welding technology has advanced, the cooling rates that can be achieved have increased, leading to inaccuracies in the predictions made with constitution diagrams.

As welding technology has advanced, the cooling rates that can be achieved have increased, leading to inaccuracies in the predictions made with constitution diagrams. Figure 1 shows two compositions that the Schaeffler diagram predicts should have similar microstructural behaviors but, clearly, do not [1]. The Schaeffler diagram shown in Figure 2 predicts that compositions 1 and 2 should have the same FA solidification mode. Multiple attempts have been made to correct these diagrams and eliminate the inaccuracies; however, most attempts are limited by the available experimental data. The objective of the current study is to develop a diagram to better predict the behavior of stainless steel welds under high cooling rate conditions by combining published experimental data with thermodynamic simulations.1174Figure 1. Micrographs of two compositions Schaeffler diagram predicts should have similar microstructures. Cr/Ni eq composition 1 is 1.51 and Cr/Ni eq composition 2 is 1.58.

Micrographs of two compositions Schaeffler diagram predicts should have similar microstructures. Cr/Ni eq composition 1 is 1.51 and Cr/Ni eq composition 2 is 1.58. Samples clearly respond differently to the etching solution and reveal significantly different features. Left image from Reference 2. 60/40 nitric acid and water electrolytic etchant for 30 seconds.Figure 2. Schaeffler diagram with compositions 1 and 2, both predicted to have FA solidification mode.1175MethodologyOne of the more prominent attempts to develop a useful diagram is a microstructural mapdeveloped by Dr. John C. Lippold, shown in Figure 3, with compositions 1 and 2 plotted for reference [3]. Lippold developed this map as a first approximation based on the published data available at the time [3]. Upon reviewing the sources that Lippold used to develop the microstructural map, it is clear that the diagram could be improved to better fit the data [3-6].

Upon reviewing the sources that Lippold used to develop the microstructural map, it is clear that the diagram could be improved to better fit the data [3-6]. Lippold used the assumption that, at low solidification rates, solidification should be dominated by the composition, not by solidification rate, and he uses this to justify having vertical boundary lines at the lower solidification rates. However, when the experimental data is plotted with the assumed line positions, the data shows a significant deviation from the assumed line positions.Figure 3. Microstructural map originally developed by Lippold [3]. Assumes weld microstructures not strongly dependent on solidification rate at lower solidification rates, indicated with vertical lines in those sections. Compositions 1 and 2 both predicted to have FA solidification mode.The cited data sources for Lippold’s microstructural map are papers written by Suutala,David et al, and a Ph.D. thesis by Elmer.

Compositions 1 and 2 both predicted to have FA solidification mode.The cited data sources for Lippold’s microstructural map are papers written by Suutala,David et al, and a Ph.D. thesis by Elmer. Each of the three papers used a different method for calculating Cr and Ni equivalency values, but also reported their full chemistries, so all values were converted to match Lippold’s method. The paper by Suutala summarizes its microstructural data with a table, and only considers three unique microstructures, labeled as austenitic-ferritic (AF), ferritic-austenitic (FA), and AF+FA, which is a combination of the previous two [4]. Suutala considered Cr/Ni values from 1.38 to 1.61 and welding speeds, which are assumed equivalent to solidification rates, from 0.416mm/sec to 13.33mm/sec.

Suutala considered Cr/Ni values from 1.38 to 1.61 and welding speeds, which are assumed equivalent to solidification rates, from 0.416mm/sec to 13.33mm/sec. Using Suutala’s data, it is clear that Lippold’s microstructural map has a poor fit with the boundary between the AF and FA regions, and, as such, modifications can be made to better match Suutala’s data [3-4]. David et al also summarized their data using a table and three unique phases. The phases used by David et al1176are fully austenitic (A), austenitic and ferritic, and fully ferritic (F) [5]. These phase designations make it difficult to distinguish between the FA and AF regions, but provide a clearer picture of the transition points for the fully austenitic and fully ferritic regions. According to David et al’s data, the boundary line between the A and AF regions could be altered to better represent the data, though the data is sparse in that region.

According to David et al’s data, the boundary line between the A and AF regions could be altered to better represent the data, though the data is sparse in that region. The data from all the papers is also sparse in the regions above approximately 1.6 for Cr/Ni equivalency and approximately 30mm/sec [3-6]. Elmer’s thesis provided additional points for this region [6]. Elmer’s thesis also provided data as to the upper boundary of the AF and FA regions, though there is ambiguity in the exact values. The points shown in Figure 4, that are credited to Elmer’s thesis, are the average of the possible range of values he presented [6]. Figure 4 demonstrates the fit between Lippold’s microstructural map and the data from the three papers.Figure 4. Lippold’s microstructural map plotted with data. Map’s fit has room for improvement. Data points presented taken from three source papers and converted to match dimensions of Lippold’s map.

Lippold’s microstructural map plotted with data. Map’s fit has room for improvement. Data points presented taken from three source papers and converted to match dimensions of Lippold’s map. Points attributed to Elmer are average values of ranges and are data points related to solidification mode transition boundaries. Elmer’s data is plotted in this manner for clarity.In the attempts to develop a more accurate microstructural map, the experimental data isoften scattered and unevenly spaced, requiring interpolation to fill in any gray areas [4-6]. It is possible to reduce these gaps in the data through the use of predictive microstructural simulations. Such simulations are possible using the Dictra module of Thermo-Calc. In order to approximate different cooling rates, the user needs to modify the physical size of the system considered and set temperature as a function of time.

In order to approximate different cooling rates, the user needs to modify the physical size of the system considered and set temperature as a function of time. The user can run solidification simulations that show when each new phase enters the system at any number of intermediate cooling rates and any desired composition.1177Once the microstructural boundaries are more clearly defined and better understood, further modifications can be made to the microstructural map to allow for the prediction of regions at high risk for failure. Such a diagram was developed for use with Schaeffler diagrams and is shown in Figure 5. The areas of concern shown in the diagram in Figure 5 are hot cracking, martensitic cracking, and sigma phase embrittlement. Due to the suppression of martensite formation that is often observed in stainless steels, the region prone to failure will likely be small, if present at all [1].

Due to the suppression of martensite formation that is often observed in stainless steels, the region prone to failure will likely be small, if present at all [1]. Sigma phase is also unlikely to be a major concern during rapid solidification, as the material cools too rapidly for it to form during solidification [7]. The hot cracking region will likely be the most significant threat, as rapid cooling and more fully austenitic microstructures tend to promote its occurrence. A more accurate map may also allow for better in-process control of the microstructure as the cooling rate could be related back to the various building parameters.Figure 5. A modified Schaeffler diagram showing regions in which welds have a high risk of failure.

A modified Schaeffler diagram showing regions in which welds have a high risk of failure. Such regions could also be overlaid on a microstructural map, providing insight on their variation with solidification rate as well as composition [8].Results and DiscussionAs shown in Figure 4, significant improvement can be made to the fit of themicrostructural map. The microstructural map predicted that both of the compositions presented in Figure 1 would most likely have microstructures that fell within the FA region. However, it appears that Composition 1 actually fell in the fully austenitic, or A, region due to the apparent microstructure. A map more accurate to both new and published data could provide a potential explanation for why the two compositions have different microstructural behavior. A key indicator of this difference is the response to the electrolytic nitric acid etchant. The etchant will preferentially attack ferrite over austenite.

A key indicator of this difference is the response to the electrolytic nitric acid etchant. The etchant will preferentially attack ferrite over austenite. Composition 1 shows few signs of having been etched, while Composition 2 shows extensive signs of etching, indicating a difference in their solidified microstructures [6].1178A more accurate microstructural map based on both the original data and the two compositions presented in this study, would still have weaknesses present in the diagram. The biggest issue is the lack of comprehensive data for the regions above 1.6 Cr/Ni equivalency and above a solidification rate of 30mm/sec. Elmer appears to be the most comprehensive source in that region, though that study only provides a limited number of data points for that region [6]. Another major weakness with the diagram is the approximate nature of its boundary lines [3]. In all the studies reviewed, there are several places where there is no data.

Another major weakness with the diagram is the approximate nature of its boundary lines [3]. In all the studies reviewed, there are several places where there is no data. For example, none of the studies presented any data for the microstructures present at compositions with a Cr/Ni equivalency between 1.79 and 1.91, despite the uncertainty as to the location of the fully ferritic transition [4-6]. The proposed simulations should be able to minimize these weaknesses, but would need to be confirmed with more comprehensive testing.ConclusionWhen used to predict the microstructure of rapidly solidified materials, constitution diagrams, such as the Schaeffler diagram, have been shown to be inaccurate. Several studies have attempted to develop diagrams to correct the inaccuracies of the Schaeffler diagram; one of the more prominent of which is the microstructural map developed by Lippold [3].

Several studies have attempted to develop diagrams to correct the inaccuracies of the Schaeffler diagram; one of the more prominent of which is the microstructural map developed by Lippold [3]. However, even these diagrams have weaknesses and could be improved in order to better predict the behavior of rapidly solidifying materials, such as additively manufactured stainless steels [3]. A review was conducted of the studies whose data Lippold used to develop the microstructural map and it was found that, according to that data, Lippold’s map has a poor fit with the experimental data [4-6]. Lippold had assumed that a weld’s microstructure was not strongly dependent on solidification rate at low solidification rates, which may have caused the diagram to poorly match the data used to create it [3]. As such, there is significant room for improvement of the microstructural map, which may be possible through the use of simulations.

As such, there is significant room for improvement of the microstructural map, which may be possible through the use of simulations. The simulations can be used to determine at what temperature each new phase enters the system, while varying the cooling rate and the composition. These simulations would allow for a much finer point distribution and, by extension, a potentially more accurate microstructural map, which could then be used to predict potential failure mechanism or to enable better in-process control of microstructure.AcknowledgmentsThis work was funded by Honeywell Federal Manufacturing & Technologies underContract No. DE-NA0002839 with the U.S. Department of Energy.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDEVELOPMENT OF AN EXPERIMENTAL LASER SINTERING MACHINE TO PROCESS NEW MATERIALS LIKE NYLON 6 Johannes Lohn1,2, Christina Kummert1, Hans-Joachim Schmid1 1Direct Manufacturing Research Center (DMRC) and Particle Technology Group (PVT), University of Paderborn, Germany 2PROTIQ GmbH, Blomberg, GermanyContact: jlohn@protiq.comAbstractSelective Laser Sintering (SLS) is an Additive Manufacturing technology which allows the production of functional polymer parts. Conventionally, Nylon 12 (PA 12), Polyamide 11 (PA 11), glass- or aluminum filled materials are used. Those materials do not always meet the requirements for direct production of serial parts by laser sintering.

Those materials do not always meet the requirements for direct production of serial parts by laser sintering. For the so called “Direct Manufacturing” of high quality, functional parts, the laser sintering process needs to be further developed and the choice of materials and needs to be expanded. During this research, a laser sintering machine for material qualification has been built up. The advantages are an optimized software solution, an innovative optical system with an adjustable laserspot, an alternative powder coating system and an improved temperature control. The functionality of the test equipment is proved with the standard material PA2200 and the new laser-sintering-material, Polyamide 6X (PA6X) is investigated. The required process parameters for processing PA6X are derived and the mechanical properties are determined by tensile tests.IntroductionSelective laser sintering (SLS) allows the direct production of functional plastic components.

A digital 3D model is cut into layers and built up by adding material. For this purpose, a thin layer of the plastic powder is applied and preheated and the powder bed surface is heated up to a temperature just below the melting point. Subsequently, a CO2 laser selectively exposures the layer-specific cross-section of the part. The plastic powder is locally melted by the laser. This iterative process of powder-coating, pre-heating and laser exposure is repeated until the part is finished. After the construction process, the additively manufactured parts are surrounded by unprocessed powder and the parts, that are not yet dimensionally stable, cool down slowly. After this cooling phase the parts can be removed from the so-called powder cake. The unmolten support powder can be sieved and reused for the next build-up process. State of the art is processing of Polyamide 12 (PA12), Polyamide 11 (PA11) and aluminum- or glass-filled compounds.

The unmolten support powder can be sieved and reused for the next build-up process. State of the art is processing of Polyamide 12 (PA12), Polyamide 11 (PA11) and aluminum- or glass-filled compounds. The use of the laser sintering process for direct production requires the qualification of engineering plastics that meet the requirements of the end products. The material PA 6 is for example known from injection molding. With the aim of substitution of injection molding by laser sintering, PA 6 needs to be qualified for laser-sintering. The current laser sintering-machines of leading manufacturers, like EOS and 3D-Systems, do not always meet the requirements to process technologically significant materials such as PA 6, because of limited processing temperature and unprecise temperature control. [SCH15]With the aim of processing new materials by SLS, an experimental SLS machine for material development was engineered and the new powder material Vestosint Z2657 PA6X was processed.

[SCH15]With the aim of processing new materials by SLS, an experimental SLS machine for material development was engineered and the new powder material Vestosint Z2657 PA6X was processed. The paper is structured as follows: First, the state of the art is described and the need557for research is clarified. After that the developed laser sintering machine, PROTIQ HT LS, is presented. Using the known material PA2200, the functionality of the high performance SLS machine is proven. Finally, the new powder material PA6X is tested, process-parameters are identified and mechanical properties are determined with tensile tests.State of the ArtThe market-leading machine manufacturers, EOS and 3D systems, use different systems for powder coating. The powder is either applied with a blade, a cassette with two blades or a counter-rotating roller [BD89][WMP10]. Conventionally used powders are optimized for selective laser sintering and characterized by a good flow behavior.

Conventionally used powders are optimized for selective laser sintering and characterized by a good flow behavior. These materials can be applied properly with the existing coating systems and offer the basis for the following process steps: pre-heating and exposure. For material development, the production of small powder quantities is important. For this purpose, polymer granules can be crushed to powder by cryogenic grinding [GHT10]. However, powder application tests in conventional laser sintering systems have shown that cryogenically pulverized powder materials cannot always be applied uniformly, smoothly and dense. Bourrel presents the idea of an alternative coating solution with a combination of a blade and a roller, to realize a high powder density in the SLS process. First the powder is applied by a blade and as a second step compacted by a rotating roller [BW14].

First the powder is applied by a blade and as a second step compacted by a rotating roller [BW14]. None of the aforementioned solutions alow processing of grinded powders with poor free-flowing properties.After powder coating, the powder surface is heated. Radiant heaters warm up the powder to a temperature just below the melting temperature. Subsequent to the application of the new powder layer, a CO2 laser selectively fuses the new powder coating with the former layers that are still molten. The molten phase is maintained by continuous surface heating while the isothermal but non-exposed powder is still solid. If the adjusted preheating temperature is too high, the entire powder cakes. If the preheating temperature is too low, the melt solidifies and shrinkage occurs. Shrinkage results in part deformation and the so-called “curling” can lead to coating errors and process break-off.

If the preheating temperature is too low, the melt solidifies and shrinkage occurs. Shrinkage results in part deformation and the so-called “curling” can lead to coating errors and process break-off. The permissible temperature range between sintering of the powder cake and the curling is material-specific. A precise temperature control is essential for processing sensitive materials [GEb07] [Sch15]. Wegner measured the temperature distribution on the powder bed surface of conventional laser sintering machines. As an example the offset-temperature of an EOS P100 is 8.7 ° C [WW13]. The inconsistent temperature distribution on the powder bed surface leads to position-dependant part properties [Dre16]. The maximum adjustable preheating temperature in the machine series EOS P3 and P1 is < 200 ° C. For this reason processing of materials like PA6, which has a melting point over 200°C, is not possible without machine modifications.

For this reason processing of materials like PA6, which has a melting point over 200°C, is not possible without machine modifications. Besides the uniform preheating, the total quantity of heat, consisting of preheating warmth and laser exposure warmth, is process-relevant. The laser energy input can be estimated by the laser energy density [MH08]:and𝑃𝑃𝑃𝑃𝐸𝐸 =𝐸𝐸𝐸𝐸 =𝑣𝑣∗ℎ𝑣𝑣∗ℎ∗𝑡𝑡The energy density (E) in J / mm² is typically defined as a function of the laser power (P), the laser spot speed (v) and the laser path spacing (h). Since the precise melt pool geometry and penetration depth is typically not known, the layer thickness (t) is used to calculate the volume energy density (EV) in J / mm³. In order to identify the material-specific energy density by sintering tests, a free558choice of the mentioned parameters is required.

In order to identify the material-specific energy density by sintering tests, a free558choice of the mentioned parameters is required. The SLS-systems, that are currently available on the market, do not offer the possibility to adjust all process relevant parameters.Conventionally for laser-exposure an optical laser system with a fixed spot diameter is used. With the aim of direct production by SLS, the speed of the SLS-process gains importance. In order to decrease exposure times, the scanning speed has been steadily increased: The first machine generation, DTM Sinterstation 2000 from the year 1992, worked with a maximum scanning speed of 1.3 m / s and the current P396 from EOS exposures with a scanning speed of 6 m / s [DTM93] [EOS16]. Proportional to the scanning speed, the laser power was increased in order to ensure the material-specific, required energy input.

Proportional to the scanning speed, the laser power was increased in order to ensure the material-specific, required energy input. Due to high laser powers, a Gaussian energy density distribution over the laser beam diameter and a low heat conduction in the plastic powder, energy peaks on the powder surface result. As a result of this energy concentration, the material can be overheated locally and degrades. Drexler detected, that material properties do not only depend on the laser energy density, but also on the interaction time of laser and powder. [Dre16] [DLD15] For those reasons, the laser energy input with expanded laser spot is investigated during this research in order to achieve a high surface exposing speed with a simultaneously low scanning speed.

The goal is a gentle, fast and homogenous energy input into the powder.Experimental set-upIn order to qualify new materials for the laser sintering process, a SLS test equipment has been developed at the University of Paderborn in collaboration with the company PROTIQ GmbH. An overview of the new SLS machine is shown in Figure 1. The maximum build space is 200 x 250 x 300 mm³. To save powder material, it is optionally possible to work with a reduced build chamber with the dimensions of 100 × 100 × 100 mm³.

The powder can be coated with the conventional solutions blade or roller and with an innovative powder application system for grinded powders with poor free-flowing properties.Laser Sintering machine for material developmentMaximum building temperature = 400°C Improved temperature control ∆T < 1.5 K • Adjustable laser spot 0.23 – 2 mm • enables high resolution and high speed exposure New slicing and machine software with open parameters and new exposure strategies Process monitoring Inert gas atmosphere with < 0,1% O2 Innovative coating system for grinded powders in three steps: Dispersing, Levelling, Compacting• •12b42a3Figure 1:SLS test equipment for material qualification (Developed at Paderborn University in collaboration with the PROTIQ GmbH)559This newly developed coater operates in three steps displayed in Figure 1: The process step Dispersing (1) dissolves existing agglomerates and improves the flow behavior.

Poorly pourable powder is rubbed mechanically through a sieve with a mesh of 150 µm. Vibrations can also be introduced to prevent the sieve from plugging. A doctor blade (2a) or a counter-rotating roll (2b) distributes the powder uniformly over the powder bed. Finally, a roller compresses the powder to the required powder density (3). In order to avoid an entrainment of oxygen and a cooling of the powder surface by the new powder layer, the powder is pre-heated in the supply container and stored under nitrogen atmosphere. A residual oxygen content of <0.1% is ensured.

A residual oxygen content of <0.1% is ensured. In order to enable an efficient and highly precise heat transport, the emission behavior of conventional and alternative beam sources was considered and compared to the absorption behavior of polyamides.Figure 2:Selection of convenient radiant heaters for efficient heating of nylon materials [Bac16] [Her16] [Tru06]It becomes clear that the emission maximum of ceramic heaters coincides with the wavelength of the laser and the absorption maximum of polyamides. In contrast, the emission maximum of conventionally used near infrared heaters is arranged in the short-wave infrared area and a considerable portion of the heat radiation is not absorbed by the powder surface. The partial reflection of the heat radiation and the rejected heat of the powder surface and the radiant heaters result in unintended warming of the entire installation space.

The partial reflection of the heat radiation and the rejected heat of the powder surface and the radiant heaters result in unintended warming of the entire installation space. As a result 12 infrared ceramic heaters are integrated into the developed machine and 12 individual heating zones are realized. The temperature distribution on the powder surface can be measured with an infrared camera and with a pyrometer. The pyrometer can be operated with a static measuring point as well as with a movable measuring spot. In order to move the measuring spot, the pyrometer is integrated into the beam path of the laser system by a beam splitter. The pyrometer and the thermal camera are protected against scattered radiation of the CO2 laser (10.6 µm) by optical filters. The pyrometer, Kleiber Infrared 660/5, measures the wavelength in a range from 4.7 - 5.2 μm. The measuring range of temperature is 30-500 ° C with a measurement uncertainty of <0.1%.

The pyrometer, Kleiber Infrared 660/5, measures the wavelength in a range from 4.7 - 5.2 μm. The measuring range of temperature is 30-500 ° C with a measurement uncertainty of <0.1%. [Kle13] The thermal imaging camera, Optris PI450 GA7, measures the wavelength of 7.9 μm. The measuring range is 200 ° C to 1500 ° C. In the extended measuring range from 0 ° C to 250 ° C it can be measured with a560reduced absolute accuracy. In order to allow an accurate measurement at lower temperatures, the pyrometer is added as a reference. [Opt16] Furthermore the position of the heating elements can be selected freely. By separately activating each individual heating element, the specific, effective range can be identified and assigned to an individual measuring field of the thermal camera. The size and position of the measuring fields is freely configurable in the machine software. The measuring fields are marked with black frames in Figure 2.

The size and position of the measuring fields is freely configurable in the machine software. The measuring fields are marked with black frames in Figure 2. With a sampling rate of 80 Hz, the average value is calculated from each measuring field and sent to the closed-loop control. The twelve-zone control ensures a uniform powder surface temperature with an accuracy of +/- 1.5 K. The maximum temperature is 350 °C.As explained before, the build up speed in the SLS process is getting more and more important with the increasing use of SLS for series production. An optical system, called Variospot, was therefore integrated into the machine. The Variospot allows the laser spot diameter to be freely adjusted within a range of 0,23 – 2 mm for each line during the exposure [Sca17]. In addition to the variable spot diameter, the focus shift is compensated and a F-Teta Lens is no longer required.

In addition to the variable spot diameter, the focus shift is compensated and a F-Teta Lens is no longer required. The use of the adjustable laser spot offers new exposure possibilities, which are illustrated in Figure 3.dl,1 :laser spot diameter in contour lineds,1 :melt pool diameter in contour line (depends on exposure parameter and material)dl,2 :laser spot diameter during area exposureds,2 :melt pool diameter during area exposure (depends on exposure parameter and material)h :distance between filling-linesBO :Beam offset between part contour and laser traceFigure 3:New exposure strategies with variable laser spot diameterDue to the variable laser spot, complex geometries with thin walls and corners, can be exposed with a small spot diameter and low laser power. In case of large areas exposure is carried out with a large spot diameter, large hatch distance and high laser power.

In case of large areas exposure is carried out with a large spot diameter, large hatch distance and high laser power. This exposure strategy allows the production of filigree parts at high build-up rates. The scanning speed and the associated scanner delays are reduced and lead to faster exposure times in total. The developed slicing and machine software enables the automatic application of new exposure strategies to CAD models.When processing new materials, dust formation and outgassing can appear during theprocess. The optics are protected by a zinc-selenium laser window, which can be cleaned during the process by means of a quick-cleaning system. Additionally the laser-window is continuously cooled and protected by an inert gas curtain.561Experimental DesignFirst, the function of the test setup is validated. For this purpose, the required process parameters for the standard material PA2200 are determined and compared to results from literature.

For this purpose, the required process parameters for the standard material PA2200 are determined and compared to results from literature. Next, the new laser sintering powder Vestosint PA6X is considered. As a basis for the subsequent sintering experiments, the particle size distribution, the particle morphology and the thermal material behavior are analyzed. Finally, the material behavior of PA6X is investigated in the sintering process. For this purpose, a uniform powder coating is attempted at first. Subsequently, the process window, determined in the first series of tests, is checked by sintering tests and the building temperature is identified. Based on this, the laser exposure is examined and the required laser energy density is determined.

Based on this, the laser exposure is examined and the required laser energy density is determined. The aim is to achieve high mechanical characteristic values and a robust SLS-process.ResultsIn order to test the functionality and stability of the experimental SLS machine, the standard material PA2200 is processed and the results are compared with literature values. For this purpose, tensile specimens with different exposure settings are manufactured and tested in order to identify the required laser energy density. For the tensile test, five tensile bars are placed in a plane with identical exposure settings. Based on sintering tests, a building temperature of 176 ° C is identified. In order to vary the laser energy input, the laser power is variable. The exposure speed and distance between filling lines (hatch distance) is kept constant. Table 1 shows the exposure parameters.

The correlation between the tensile modulus of elasticity and the selected energy density is shown on the left. As explained by Grießbach, an energy plateau that leads to robust mechanical properties is characteristic for the material PA2200 [Gri12]. This energy plateau is reached from a volume energy density of 0.25 J / mm³. Wegner describes an energy plateau of 0.3-0.35 J / mm³ for PA2200 [WW13]. In our investigations, an energy plateau in the range of approximately 0.22-0.40 J / mm³ was identified. For subsequent experiments a volume-energy-density of EV = 0.3 J / mm³ is used, to identify the mechanical characteristics by tensile tests.562Figure 4:Results of parameter study with PA2200In Figure 4 the function of the developed SLS machine is validated by comparing the results of the parameter study with values from the product data sheet of PA2200. The table shows that the exposure parameters of the experimental machine lead to similar mechanical properties.

The table shows that the exposure parameters of the experimental machine lead to similar mechanical properties. A higher modulus of elasticity and an increased tensile strength are noticeable in combination with a reduced tensile elongation. This effect may be due to the testing without conditioning. With the absorption of water, typically the modulus of elasticity and tensile strength decrease while the values of the tensile elongation increase.In the next step tests with the new SLS powder material Evonik PA6X are performed. Figure 5 compares the particle morphology of the investigated PA2200 and PA 6X.

Figure 5 compares the particle morphology of the investigated PA2200 and PA 6X. The chemically precipitated materials, PA 2200 and PA 6X, are characterized by rounded particles with high sphericity.% n ie m u l o v e v i t a l e r20.0010.00PA2200PA6X PA 6130.00PA2200PA6X250 µm10.00particle size (log)100.00Figure 5:Particle morphology and particle size distribution of the investigated materials PA2200 and PA6XIn addition to the particle morphology, the particle size distribution has a considerable influence on the flow and application behavior during powder application in the sintering process. If the particle size distribution is wider, a poorer flowability results, even if the median value of the particle size is equal [Sch14]. The particle size distribution in Figure 5 was determined by means of laser diffraction with the measuring device Mastersizer 2000 from Malvern Instruments.

The particle size distribution in Figure 5 was determined by means of laser diffraction with the measuring device Mastersizer 2000 from Malvern Instruments. Based on the results a high powder flowability and a smooth powder surface after coating can be foreseen.563Within the scope of the powder characterization, the material-specific thermal behavior is also examined by means of dynamic differential calorimetry. The thermal behavior of the materials is tested according to DINISO / DIN 11357. A heating and cooling rate of 10 K / s is selected. Based on the results in Figure 6 the required preheating temperature for PA6X can be estimated to be in the range of 185 – 200 °C.Figure 6:Thermal behavior of PA2200 and PA6X – Measured with dynamic differential calorimetry (DINISO / DIS 11357)Analogously to the test procedure with PA2200, the new material PA6X is processed. Table 2 shows investigated exposure parameters.

The determined building temperature is 192 ° C. The sintering process is stable and robust in a range of temperature from 188 °C up to 195 °C. The white color of PA6X is similar to sintered parts out of the material PA2200. Very filigree models can be realized and the residual powder does not tend to cake. With the minimum laser spot of 0.23 mm a wall thickness of 0.23 mm can be achieved. Figure 8 shows that the required energy plateau for the material PA6X is reached at a laser volume energy density level of 0.35 J / mm³. From an energy density of > 0.4 J / mm³, the tensile strength begins to decrease, because of degradation.564Figure 7:Tensile strength as a function of Laser Energy Density during exposureConclusions & OutlookAn SLS system was developed for the qualification of new LS materials.

The main benefits of the SLS test equipment are an optimized temperature management, an innovative powder application for grinded powder materials and a freely adjustable laser spot diameter. All machine and process parameters can be selected freely inside the developed slicing and machine software. The function of the system technology was validated with the known material PA2200 from EOS. For this purpose tensile specimens were built and the volume energy density of the laser exposure was varied. The necessary energy level for a robust process management was identified. The obtained mechanical characteristics correspond to the data in the material data sheet of the manufacturer EOS for PA2200.In addition, the new SLS material PA6X was processed. PA6X is characterized by good processability. The highest tensile strength of 58 MPa was achieved with a laser power density of 0.37 J / mm³.

PA6X is characterized by good processability. The highest tensile strength of 58 MPa was achieved with a laser power density of 0.37 J / mm³. The material PA6X stands out against the standard material PA2200 by a higher temperature resistance and a high elongation at break > 25 %.In future work, the effect of different exposure strategies will be investigated and a statistical tests will be used to confirm the gathered mechanical properties. In particular, the influence of a variable laser diameter on the material properties has not yet been conclusively investigated. In order to make comparison with existing laser sintering systems with a constant laser spot diameter possible, the area exposure rate with the unit mm² / s can be an interesting characteristic parameter. The laser exposure time per unit of area at identical surface speed might be increased by the exposure with an expanded laser spot.

The laser exposure time per unit of area at identical surface speed might be increased by the exposure with an expanded laser spot. It is known from injection molding that, compared to polyamide 12, polyamide 6 absorbs significantly more moisture after the manufacturing process, thereby mechanical properties and dimensions can change. For this reason water intake and its influence on the mechanical characteristics should be investigated. So far the thermal behavior has been determined by means of DSC analysis. For the design of serial products, mechanical properties will be determined as a function of the application temperature.565AcknowledgementThe authors want to thank the Phoenix Contact foundation as well as all partners and colleagues of the Direct Manufacturing Research Center.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDEVELOPMENT OF AN EXPERIMENTAL TEST SETUP FOR IN SITU STRAIN EVALUATION DURING SELECTIVE LASER MELTINGEckart Uhlmanna, Erwin Krohmera, Felix Hohlsteinb, Walter ReimersbaInstitute for Machine Tools and Factory Management (IWF), TU Berlin, Pascalstraße 8-9, Berlin, Germany bInstitute for Material Science and Technologies, Metallic Materials, TU Berlin, Ernst-Reuter-Platz 1, Berlin, GermanyAbstractSelective Laser Melting (SLM) is an Additive Manufacturing (AM) process which still underlies a lack of profound process understanding. This becomes obvious when deformation and crack formation can be observed in SLM parts due to residual stresses. Controlling residual stresses is therefore an important topic of recent research in AM of metals.

Controlling residual stresses is therefore an important topic of recent research in AM of metals. In order to minimize residual stresses further knowledge considering their cause and physical correlations of process parameters needs to be generated. In this paper an approach of measuring strains layer by layer during the SLM process by means of in situ X-ray diffraction is presented. For this purpose an experimental test setup is being constructed at the Technische Universität Berlin. The system requirements and operating principles are discussed in this paper. Furthermore, details of the current progress of the construction are highlighted.IntroductionIn metal based Additive Manufacturing (AM) a continuously growing market can be observed [1]. Among the variety of different AM technologies the powder bed fusion, especially the Selective Laser Melting (SLM), is the leading technology for metal parts [2].

Among the variety of different AM technologies the powder bed fusion, especially the Selective Laser Melting (SLM), is the leading technology for metal parts [2]. However, there are still some characteristic drawbacks as a result of the physical processes during SLM. High thermal gradients during the building process are a result of the locally concentrated energy input by the laser source and lead to residual stresses in the part [3]. As a consequence, the mechanical properties as well as the geometrical accuracy are negatively influenced and deformation as well as crack formation is likely to occur. In state-of-the-art machinery the parts are fused on a substrate plate either directly or in combination with support structures. The connection between substrate plate and part leads to a better heat transfer and prevents deformations of the part. Still, the residual stresses can result in delamination of the SLM parts from the substrate plate as depicted in figure 1.

Still, the residual stresses can result in delamination of the SLM parts from the substrate plate as depicted in figure 1. Furthermore, by heating the substrate plate, and by that also the part, up to 550 °C [4] the temperature gradient and thus the thermal stresses are reduced. The use of brittle materials like carbide metals [5] or ceramics with increased melting points and therefore the likeliness of crack formation is still limited and could be improved by an enhanced high- temperature preheating of more than 1600 °C [6]. Other strategies to reduce residual stresses in AM parts have been investigated such as the scanning strategy [3] and the heat treatment after the building process [7, 8].

Other strategies to reduce residual stresses in AM parts have been investigated such as the scanning strategy [3] and the heat treatment after the building process [7, 8]. Though the mechanisms leading to residual stresses during SLM basically are known, a deeper understanding of their complex correlations with the SLM process1472parameters and dependencies of magnitudes is needed for the improvement of process strategies leading to better part quality.Figure 1: Delamination of Ti6Al4V SLM parts from the substrate plateThere are a number of approaches to evaluate residual stresses in AM parts reported in the literature. The different techniques are based on the measurement of strains or part deformations that are correlated to the compensation of internal stresses. A lot of experiments have already been conducted regarding the post-process measurement of strains in SLM-built parts by means of neutron diffraction [9, 10] or X-ray diffraction [11-13].

A lot of experiments have already been conducted regarding the post-process measurement of strains in SLM-built parts by means of neutron diffraction [9, 10] or X-ray diffraction [11-13]. However, X-ray diffraction is only applicable to measure the remaining strains on the surface of the part. In addition, incorrect values may be obtained as the penetration depth of the X-rays often is comparable in magnitude to the values of surface roughness of metal AM parts [14]. While neutron diffraction is suited for deep measurements in components there are disadvantages regarding the relatively low time resolution as neutrons interact weakly with matter. Other methods like the crack compliance method [3, 15], contour method [10, 16], layer removal method [7, 17, 18] or the hole drilling method [14, 19] have been evaluated for SLM parts, but are destructive respectively semi-destructive methods and therefore have to be applied after the building process.

An approach of investigating the stresses during the process was performed by attaching strain gauges to the bottom of the build plate [20]. By that the strains could be measured layer by layer during the process in order to evaluate the residual stresses. However, with this method a spatial resolved determination of stresses within the part is not possible as the stress components in part height direction cannot be measured. New measurement methods are required to clarify relationships between process parameters, signatures and part quality and the relative sensitivities of those relationships through experiments [21].With brilliant synchrotron light sources a highly resolved spatial investigation of stress distribution and microstructural changes during dynamic production processes became possible [22] as already shown by investigations of the chip formation zone during in situ cutting experiments [23, 24].

In order to improve the understanding of the residual stress formation during SLM due to the complex interaction of the temperature gradient mechanism, subsequent thermal cycling, phase transition and recrystallization in a spatial resolved manner, a new test setup principle for in situ strain evaluation by means of synchrotron X-ray diffraction is presented in this paper.1473Operating principleWith respect to the synchrotron X-ray diffraction method there are some major requirements for the design of the SLM process chamber which have to be fulfilled for the successful application of in situ experiments during SLM to collect spatial data of the stress distribution in the part. In the first place the transmission of X-rays must not be unnecessarily interfered throughout the process chamber to ensure a preferably unfiltered measurement and short acquisition times.

In the first place the transmission of X-rays must not be unnecessarily interfered throughout the process chamber to ensure a preferably unfiltered measurement and short acquisition times. Secondly, the synchrotron beam as well as the detector is stationary, meaning that the relative movement between powder bed and laser beam spot has to be established by a moving powder bed. The working principle of the test setup is shown in figure 2a. Pursuing this working principle the measuring mode depicted in figure 2b can be enabled.Laser (x = const., z = const.)Movement of base plateSynchrotron beamSynchrotron radiationScanner opticsPowder housingΔxΔzMetall powderProcess chamberLaser2D detectorSubstrate platea)b)zyxFigure 2: a) Operating principle; b) Measurement mode of the test setupA measuring location in a horizontal distance Δx and a vertical distance Δz to the laser is defined.

Due to the stationary synchrotron beam also the laser has to maintain its position in x and z direction to keep the distances Δx and Δz constant during measurement. The relative movement between powder bed and laser in x-axis direction is enabled by an actuated base plate. Since the base plate moves in x-axis direction, the laser oscillates in y-direction between the powder housing so the complete powder layer can be exposed to the laser and synchrotron beam while keeping a constant distance between part and detector. Thus the strains are measured integrally and sufficient diffraction intensity can be measured during exposure time. Regarding the depicted operating principle in figure 2a, the strains in part width direction i.e. y-direction cannot be measured because of the two-dimensional character of this measurement technique.

y-direction cannot be measured because of the two-dimensional character of this measurement technique. By providing a possibility to rotate the test setup around its z-axis, further measurements under different incidence angles can be conducted and thus the stress state in part width direction can be deduced.Functional RequirementsConsidering the mentioned requirements due to the operating principle having an impact on the process itself, there are other requirements which have to be fulfilled on top. Owing to the available amount of space and weight limits of positioning systems in synchrotron laboratories,1474e.g. an hexapod as stated in [24], the simple utilization of conventional SLM systems is not practicable for this purpose. As a result a SLM system with an adapted process chamber has to be designed.

As a result a SLM system with an adapted process chamber has to be designed. In table 1 the functional requirements determined for the system which is currently developed at the Technische Universität Berlin are listed.Table 1:Functional requirements for the design of the in situ SLM systemRequirementSpecificationBuild platform length lb Build platform width wb Build envelope height hb Build platform heating temperature Th Laser power Pl Laser focus diameter dl Powder housing X-ray transmission I/I0 Powder housing operating temperature Top Powder bed effective x-axis stroke seff Maximum mass mmax (mounted on positioning system)140 mm3 mm10 mm400 °C400 W, continuous wave80 μm to 500 μmpreferably unfiltered1800 °C140 mm250 kgThe long term objective of this current work is to map process parameters with the resulting residual stresses and therefore contribute to the process understanding in metal AM, leading to improved process strategies.

The system should be able to produce basic parts like simple thin walls or parts with features, as listed in table 2, with physical and mechanical properties comparable to common SLM machines. Hence the experimental system is designed to work in parameter ranges known from common SLM systems. The exceptions are the possible part thickness which is limited by the attenuation of the incident X-radiation across the material and the laser scan speed which is limited to some extent by the inertia of the powder in the moving powder bed. The Laser power as well as the Laser focus diameter is in a comparable range as in the SLM 250HL machine from SLM SOLUTIONS GROUP AG, Lübeck, Germany, which operates at the Institute for Machine Tools and Factory Management. The adjustable process parameters dependent on the system are shown in table 2.

The adjustable process parameters dependent on the system are shown in table 2. Further process parameters which affect the formation of residual stresses, determined by the used metal powder or the part geometry, are also accessible for investigations on residual stresses. As the synchrotron radiation penetrates through the complete powder bed, the complete powder layer has to be molten and solidified in order to avoid incorrect measurements. This is to avoid deflection of the incident X-rays by unsolidified powder that doesn’t contribute to the effective residual stress distribution in the built part. In figure 3 the different available scan path strategies of the test setup are shown. It can be seen that the scan strategy presumably leading to the highest residual stresses, which is the x-hatching, is also the scan strategy where the most unsolidified powder is measured if the measured layer equals the currently processed layer.

Scan strategies which involve diagonal scanning or any rotational angle other than 0° or 90° will not be possible to carry out with the introduced experimental system.1475Table 2:Process parameters accessible for investigationMachine parametersPowder characteristicsPart geometry Laser power  Laser scan speed  Laser focus diameter  Hatch distance  Scan path  Layer thickness  Base plate temperature  Base plate material  Inert gas flow Material  Particle size distribution  Bulk density  Flowability  Humidity  Shape/sphericity Solid parts  Hollow parts  Overhang  Bore holes  Slots  Edges  Support structuresOne of the most challenging questions in the experimental setup design is how to avoid undesired attenuation of the synchrotron X-radiation as a result of machine components which are located in the beam path. First, the powder bed has to be elevated and surrounded by a X-ray transmissive powder housing.

First, the powder bed has to be elevated and surrounded by a X-ray transmissive powder housing. Second, the process chamber housing must be transmissive where the beam path crosses. According to the first requirement there is the need to identify a suitable material which ensures not only the X-ray transmission but also has to endure high temperatures above the melting point of common SLM materials, as for example titanium with its melting point at 1668 °C. Monocrystalline sapphire (Al2O3) was identified as a material to meet the requirements with the melting point at 2050 °C and a maximum operation temperature of 1800 °C. The theoretical attenuation coefficients of sapphire can be determined partially [25]. The curve in figure 4 is fitted by piecewise cubic Hermite interpolation and shows the correlation of the linear attenuation coefficient μ and the incident photon energy E.

The curve in figure 4 is fitted by piecewise cubic Hermite interpolation and shows the correlation of the linear attenuation coefficient μ and the incident photon energy E. Further, the attenuation is calculated exemplary for the photon energies 50 keV, 100 keV and 200 keV respectively for the transmission through a sapphire plate of 2 mm wall thickness. The obtained values show the general influence of the energy level of the incident X-radiation on the attenuation in the material. A decreasing attenuation for increasing values of the photon energy is characteristic for the relevant energy interval. In addition, a few percent of transmission is already sufficient to get satisfactory intensity results at the detector because of the high photon flux in synchrotron radiation [26].

In addition, a few percent of transmission is already sufficient to get satisfactory intensity results at the detector because of the high photon flux in synchrotron radiation [26]. Hence the transmission through the sapphire plates for high photon energies is theoretically adequate.UnidirectionalAlternatingyxyxFigure 3: Scan path strategies14761500,70keVCalculationbasis:Beer-Lambert lawI = I0∙e-μtAttenuationA = 1-e-μtMaterial:Sapphire(Al2O3)Densityρ= 3.98 g/cm³Wall thicknesst = 2.00 mm2460%12AttenuationA502501001,400,350,00Photon energyELinear attenuationcoefficientμ1cmFigure 4: Theoretical attenuation coefficients for sapphire (Al2O3) [25] and exemplary attenuation through2 mm wall thickness dependent on the incident photon energyThe applicability of sapphire plates as powder bed housing in respect of the high energy input by the laser was tested in a powder bed of Ti6Al4V material in a SLM 250HL machine.

The machine is equipped with a Nd:YAG Laser with a wavelength of λ = 1064 nm and a laser focus diameter of dl = 70 μm. The laser power was Pl = 275 W at a scanning speed of vs = 975 mm/s and a hatch distance of h = 120 μm. With a layer thickness of Δz = 50 μm the energy density can be determined to Ed = 47 J/mm³. The used sapphire plate had a thickness of t = 0.5 mm and measured 10 mm both in length l and in width w and was placed upright in the powder bed. The operating area surrounding the sapphire plate was 10 mm x 10 mm. There were n = 10 trials scanned in this area while reducing successively the distance between scanned area and sapphire plate from both sides. The scanning method is shown in figure 5a. Until the last trial no influence by the laser energy input on the sapphire plate could be seen visually. In the last trial the sapphire plate was scanned directly by the laser. After the last trial the effect of the energy input as seen in figure 5b could be noticed.

In the last trial the sapphire plate was scanned directly by the laser. After the last trial the effect of the energy input as seen in figure 5b could be noticed. The upper edge is rounded and cracks occur partially. A direct hit by the laser spot is therefore to be avoided, as the changed shape of the upper edge might influence powder recoating and X-ray diffraction measurements.10 mmPowder beda)b)2mmScanned areaSapphire plate0.5 mmnFigure 5: a) Sapphire plate scanning method with n = 10 trials; b) Sapphire plate after direct impact by the laser1477Current stage of constructionIn figure 6a the conceptual design of the in situ SLM system is shown. The gas filtration unit as well as control unit and user interface are situated separately and therefore help to meet the weight requirements. The laser security windows make sure that the incident radiation as well as the deflected radiation is able to traverse the process chamber.

The laser security windows make sure that the incident radiation as well as the deflected radiation is able to traverse the process chamber. Another laser security window on top of the chamber allows the visual monitoring of the process. For the operating mode a telescopic beam stop is mounted between scanner and process chamber to ensure safety operation. The displaceable scanner system guarantees the possibility of adapting the operating distance when mounted on the synchrotron laboratory positioning system. The adapter platform which is needed to be mounted on such a positioning system is not implemented yet. Figure 6b pictures the moving build platform design. The z-axis and x-axis of the build platform are decoupled, meaning that the powered linear bearings are not mounted on each other. This is necessary in order to keep the sapphire powder housing permanently on the same z-level. Else, the powder recoating mechanism would have to be adjustable in z-axis direction.

This is necessary in order to keep the sapphire powder housing permanently on the same z-level. Else, the powder recoating mechanism would have to be adjustable in z-axis direction. The build platform is mounted on a rail that is fixed on the actual build cylinder, which serves as actuated z-axis. A sliding contact between the sapphire plate carrier and the build platform ensures that the x-axis movement is transmitted whereas the z-axis movement of the cylinder only affects the build platform itself. The stroke of the z-axis cylinder is constricted by the build platform rail, but is enough to reach the required build envelope height.

The stroke of the z-axis cylinder is constricted by the build platform rail, but is enough to reach the required build envelope height. In the future design a powder recoating mechanism will be implemented in the in situ SLM machine.linear axissapphire platespowder bedX-axis driveLinear axisSliding contactBuild platform railx-axis drivebuild platform railb)z-axis cylindersliding contactZ-axis cylinderPowder bedSapphire platesa)Figure 6: a) Conceptual design of the in situ SLM system; b) Build platform designConclusionsIn this work the need for new time-resolved measurement techniques for residual stresses in SLM parts was outlined. Based on this need, the operating principle of a SLM system for in situ synchrotron X-ray experiments was clarified. The operating principle demands to restructure and redesign the common known SLM process chamber, as a moving powder bed has to be installed.

The operating principle demands to restructure and redesign the common known SLM process chamber, as a moving powder bed has to be installed. Further, some modifications concerning the transmittance of X-radiation throughout the process chamber have to be made. The powder bed has to be elevated above the build platform where sapphire was identified as a suitable material for the transmissive powder housing in regards of1478the attenuation of incident X-rays as well as high temperature resistance demands. Further, the current progress of construction of the in situ SLM system at the Technische Universität Berlin was illustrated.AcknowledgmentsThis paper is based on results acquired in the project DFG UH 100/207-1, which is kindly supported by the German Research Foundation (DFG).References[1] Wohlers, T.: Wohlers Report 2015: 3D printing and additive manufacturing state of theindustry; annual worldwide progress report. Wohlers Associates, Fort Collins, Colo., 2015.

E DEVELOPMENT OF AN INTEGRATED GRAPHICAL SLS PROCESS CONTROL INTERFACEGuohua Ma and Richard H. Crawford The University of Texas at AustinABSTRACTThis paper presents the systematic development of a man/machine interface for SelectiveLaser Sintering. The interface developed integrates geometry design, process data/code generation, power control and communication in a graphical environment. The objective of this interface is to provide high flexibility and robustness to researchers to meet their special needs. For example, the user can generate laser power profiles of any shape, or control the laser power by PWM or an analog voltage. The control code for the machine is automatically generated with minimum human input. Currently the interface is limited to use on a single layer machine, but can be easily modified for multi-layer machines.

Currently the interface is limited to use on a single layer machine, but can be easily modified for multi-layer machines. LabVIEW™ is employed as the developing platform.Keywords: Selective laser sintering, interface software, process control, LabVIEWINTRODUCTIONSelective Laser Sintering (SLS) is a process that employs a powdered material approach to rapid prototyping. The process begins with the deposition of a thin layer of powder, which is pre-heated to just below its melting point. A laser selectively traces the surface of the powder and sinters the material together. This process continues layer by layer until a complete part is fabricated [2].Commercial SLS workstations are produced by 3D Systems. With these machines, theuser can control the laser scanning parameters to finish basic SLS steps.

With these machines, theuser can control the laser scanning parameters to finish basic SLS steps. The goal of the software system on these commercial workstations is to limit the amount of information required from the user, while producing consistent, repeatable part quality. However, researchers in the Laboratory for Freeform Fabrication (LFF) at UT Austin often build laboratory-sized SLS workstations, and must develop control software for each of these machines. The goal of this work reported here is to develop a unified control software system with a unified interface that can be ported to every laboratory workstation with minimum modification.To support the research, the interface must be very flexible to accommodate the specialrequirements of different SLS experiments. For example, the power level as a function of time is a very important factor that will influence the quality of the final part and should be optimally designed for different materials and geometries.

Secondly, the overall SLS process should be integrated into a single user-friendly environment to reduce the configuration time. Configuration of each SLS process, such as control code generation configuration, should be flexible and simple enough to minimize required human input. Thirdly, the interface should™ LabVIEW is a product of National Instruments Corp., Austin, TX, USA.408provide the option of simulating the SLS process before execution for cost and safety considerations.To meet above requirements, a graphical user interface and necessary hardware interfaceto a General Scanning controller was developed using National Instruments’ LabVIEW™ software. Although this interface is designed specially for use in a SFF research laboratory, it can be adapted for commercial use as well.SYSTEM OVERVIEWThe interface was targeted for two different workstations in the LFF. Each workstationuses a General Scanning, Inc.

Each workstationuses a General Scanning, Inc. (Watertown, MA) DE2000 scan controller and XY3037 scan head. One workstation has a Nd:YAG laser beam, while the other uses SYNRAD’s 48 series CO2 laser beam. The interface accommodates either configuration.The structure of the interface system is shown in Figure 1. The major functions of theinterface are to accept information from the operator, generate the control code and power profile, and communicate with the SLS machine controller. The hardware consists of a parallel port (for control code download) and a National Instruments AT-MIO-16E series DAQ board.Figure 1. System structure.The main interface designed in LabVIEW 6.0 is shown in Figure 2. The menus from leftto right follow the SLS operation sequence. When the user selects one of the menus, a detail window corresponding to this menu pops up. The menus are Geometry, Toggle File, Process, Power Design, Simulation, and Run.

When the user selects one of the menus, a detail window corresponding to this menu pops up. The menus are Geometry, Toggle File, Process, Power Design, Simulation, and Run. A new operator can just follow the menus left to right to complete a part. An experienced operator can select specific menus without following the left to right sequence. Communication between the SLS workstation and the control computer takes place through the parallel port and the NI DAQ board.™ LabVIEW is a product of National Instruments Corp., Austin, TX, USA.409USER INTERFACE: GEOMETRYThe SLS software uses an off-line preprocessor to convert an input STL file to a laser toggle point file that controls the SLS workstation. However, when conducting experiments, simple geometries such as circles or squares are always used first. Thus, the GUI provides a convenient method for designing simple geometries.Figure 2. Geometry design module.Figure 2 shows the geometry design interface.

Thus, the GUI provides a convenient method for designing simple geometries.Figure 2. Geometry design module.Figure 2 shows the geometry design interface. Currently, both open curves, such as lines,circles, and spots, and areas, such as squares and circles, are supported. Each geometry input is represented as an entity; entities can be combined to form more complex geometry. The advantage of this approach is that the user can adjust the number of scan lines to change the quality or properties of the part. For example, suppose a certain experiment requires a U-shaped area, with different densities for different sub-areas of the program. Commercial SLS workstations assume the desired density of the part is constant for a given build.

Commercial SLS workstations assume the desired density of the part is constant for a given build. With our research interface, this desired outcome is readily realized.Although this module can only deal with simple geometries, it provides the “all-in-one” ability that allows researchers to design different experiments using a single software package. This module can be expanded to include a pre-processor module for STL files, if necessary.USER INTERFACE: TOGGLE FILEThe toggle file generation module is one of the most important modules of the system. . The function of this module is to rasterize the geometry information into laser power toggle point information. The toggle point information is the coordinates where the scan lines intersect the contour of the geometry. The interface to this module is shown in Figure 3.In this module, the user can choose the scanning pattern, which directly affects thequality of the part.

The interface to this module is shown in Figure 3.In this module, the user can choose the scanning pattern, which directly affects thequality of the part. The General Scanning DE2000 controller supports both unidirectional and bi-directional scanning. Thus, the toggle file module allows the user choose either a zigzag or410bi-directional rasterization. The user can also choose either the horizontal or vertical direction as the major scanning axis.Figure 3. Toggle file generation module.USER INTERFACE: PROCESSThe process module converts the toggle point information to DE2000 controllercommands. It requires four different types of information: file information, scanning mode information, part parameters, and laser scanning parameters. All of this information is translated into the format required by the controller.

All of this information is translated into the format required by the controller. The interface helps the users set the parameters, and the controller code is generated automatically based on the toggle file and the user-provided parameters. The interface is shown in Figure 4.USER INTERFACE: POWER DESIGNLaser power is a major factor that influences SLS part quality. In commercial SLS workstations, the laser power remains fixed during the entire part build. However, our research calls for more extensive control of laser power during a part build. For instance, control of laser power is very important to eliminate balling in direct SLS metal processing [1]. For different geometries and materials, a specially designed power profile may improve the quality of the part. In many cases, determining these relationships is a trial and error process.

For different geometries and materials, a specially designed power profile may improve the quality of the part. In many cases, determining these relationships is a trial and error process. Thus, the user interface provides much flexibility for the user to control the power profile of the SLS workstation as desired.Figure 5 shows the interface for specifying the power profile. The user has three options for generating the power profile: (1) create the profile from library functions (e.g., sine, triangle etc.); (2) sketch the profile; and (3) load the profile from an external file. With this module, the user can quickly create practically any desired power profile.411Figure 4. Process module interface.Figure 5. Power profile design module.As previously mentioned, two different laser systems are used in the LFF; therefore, twodifferent control systems are employed. The Nd:YAG laser uses an analog voltage to directly control the laser power.

The Nd:YAG laser uses an analog voltage to directly control the laser power. Control of the SYNRAD 48 series CO2 laser is achieved by pulse width modulation (PWM). This mode can be used to control instantaneous power by adjusting the pulse width (PWM duty cycle) at a fixed modulation frequency.In the voltage control mode, the laser power is directly proportional to a control voltage. This conversion is accomplished by directly sending the power profile to the controller through the DAC port on the NI DAQ board.412Under PWM control mode, the power is controlled by a fixed frequency PWM signal.The 48 series laser beam requires a 1 µs “tickle” pulse delivered at a 5kHz clock frequency from the controller. The standard PWM frequency is 5kHz, which has a period of 200 µs, as shown in Figure 6 [3]. A counter in the NI board is used to generate the PWM pulse train and the tickle signal [4].Figure 6.

The standard PWM frequency is 5kHz, which has a period of 200 µs, as shown in Figure 6 [3]. A counter in the NI board is used to generate the PWM pulse train and the tickle signal [4].Figure 6. PWM control durationThe relationship between the output power and the PWM pulse duty cycle is measuredexperimentally. As shown in Figure 7, when the duty cycle is less than 60%, the relationship between power and duty cycle is very linear, and becomes somewhat nonlinear when the duty cycle is higher than 60%. The duty cycle is generated based on required power from the data pictured in Figure 7.Laser Calibration - 2/11/200230.025.0)W20.0(r e w o P15.010.05.00.0020406080100Duty cycle (%)Figure 7. Calibration of laser power and duty cycle for PWM mode.413USER INTERFACE: PROCESS SIMULATIONThe process simulation module is still under development. The initial design is shown inFigure 8.

Calibration of laser power and duty cycle for PWM mode.413USER INTERFACE: PROCESS SIMULATIONThe process simulation module is still under development. The initial design is shown inFigure 8. In this module, the process information, including the layer profile, laser power and other process information is displayed in a single window. The power versus time and laser position are displayed so the user can assess their effects before the real part is fabricated.Figure 8. Simulation module.USER INTERFACE: RUNAs shown in Figure 9, all information generated in previous modules is integrated in the run module. The geometry, power profile, power control, and other information are displayed. The power control mode is set in this module. The working sequence of this module is shown in Figure 10.Figure 9. Run module.414Figure-10. Operation sequence of run module.The user selects power control mode and then click on the START button.

Run module.414Figure-10. Operation sequence of run module.The user selects power control mode and then click on the START button. The toggle point file is then downloaded to the SLS Controller. Once the download is complete, the SLS Controller sends a trigger signal back to the GUI and starts to execute the file. The trigger signal activates the output of the power control signal. Thus the motion of the laser deflection mirrors is synchronized with the power profile. The temperature monitoring function is still under construction.CONCLUSION AND FURTHER DEVELOPMENTThis paper describes a man/machine interface for controlling a research SLS workstation.The interface integrates geometry design, process data/code generation, power control, and communication with the workstation in a graphical environment.

The main objectives of this interface are to provide high flexibility and integrated information display to meet the special needs of the researchers who will use the SLS workstation for a variety of experiments.Further improvements to the GUI are underway. More extensive geometric designcapability is being added, including a STL file interface. Real-time monitoring of the process temperature, which will give researchers the opportunity to study the relationship between temperature and laser power, part quality, etc., is also being developed. In current SLS workstations, only open-loop control is employed, and the laser power cannot be adjusted during the part build. Employing closed loop control of laser power based on the operating temperature is an important goal of future development.REFERENCES1.

Employing closed loop control of laser power based on the operating temperature is an important goal of future development.REFERENCES1. Das, Suman “Direct Selective Laser Sintering of High Performance Metals – Machine Design, Process Development and Process Control”, Dissertation, University of Texas at Austin, 1998.2. Beaman, J. J., Barlow, J. W., Bourell, D. L., Crawford, R. H., Marcus, H. L., and McAlea, K. P., Solid Freeform Fabrication: A New Direction in Manufacturing, Kluwer Academic Publishers, 1997.3. SYNRAD, “SYNRAD 48 series CO2 laser operation manual”. J version http://www.synrad.com/Products/manuals.htm.4. National Instruments, AT-MIO-16E series DAQ board manual415.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDevelopment of automatic smoothing station based on solvent vapour attack for low cost 3D PrintersProf. Marlon Wesley Machado Cunico, Ph.D. , marloncunico@yahoo.com.br Universidade de São Paulo campus São Carlos Concep3d Pesquisas CientíficasPatrick Medeiros Cavalheiro Concep3d Pesquisas CientíficasProf. Jonas de Carvalho, Ph.D. Universidade de São Paulo campus São CarlosAbstract: Along the last years, 3d printing has been playing a new and important role in several market segments. As consequence, finishing methods have been developed and applied in order to improve surface roughness and mechanical strength. One of these methods is the solvent vapour attack. Nonetheless, this process is still manual and might lead to object deformation or structural damages.

One of these methods is the solvent vapour attack. Nonetheless, this process is still manual and might lead to object deformation or structural damages. For that reason, the main goal of this work is to present a new approach that was implemented in automatic smoothing station. In this new approach, a close-looping control system identifies the vapour attack level in addition to controlling drying time and number of times that cycle is repeated. By the end, this proposal was identified to advances in 3d printing field, being a next step for domestic and distribuited manufacturing.Key words: Additive manufacturing, Smoothing process, dimensional evaluation1. INTRODUCTIONAdditive manufacturing technologies have remarkably grown along the last years. Additionally, fused filament fabrication or low cost fused deposition modelling has also achieved an extraordinary advance with respect to precision and finishing [1].

Additionally, fused filament fabrication or low cost fused deposition modelling has also achieved an extraordinary advance with respect to precision and finishing [1]. Nevertheless, objects which are obtained by such methods use to result in anisotropic mechanical behaviour besides poor surface roughness. For that reason, low cost equipment is mostly restricted to hobby and prototypes applications.In order to solve or reduce these issues, Priedeman and Smith [2] introduced the surface post- process based on vapour solvent attack, whose patent owner is Stratasys. This post-process is also known as smoothing process. In general, the surface finishing of the object is provided by a vaporised solvent treatment.

This post-process is also known as smoothing process. In general, the surface finishing of the object is provided by a vaporised solvent treatment. Other works studied different kinds of smoothing techniques, so that we can classify these processes in: a) liquid immersion [3, 4], b) cold vapour [4, 5]; c) cold atomizer vapour [6]; d) hot vapour [1, 7, 8].In spite of that, these kind of processes are still hand craft and might imply in object geometrical distortion in addition to object damage[1]. In addition, different object shapes make the process control harder. Therefore, the main goal of this work is to introduce a new method that controls the vapour attack process in an automatic way.In this new method, we developed and fabricated an automatic smoothing station that controls the severity of hot vapour attack, level of drying, temperature, pressure and solvent absorption rate in a1982close-loop system.

The simplified schematic of this concept is presented in Figure 1, where the basic diagram block of the close-loop control system is also exposed.Figure 1 –Schematic of automatic smoothing station and block diagram of close-loop control systemIn order to evaluate this proposal, we analysed the object roughness as a function of solvent absorption ratio, drying ratio and number of cycles. In this analysis, besides identifying the transfer function of system, we evaluated the efficiency of On-Off control system with target and minimum threshold.It is important to note that the load cell is placed outside of smoothing chamber, and the condensated liquid on the platform was measured and removed from the threshold, compensating the effect of liquid on surface during the evaporation and drying stages.By the end, we identified the main implications and potential application of this new method for small and medium business and specialised market segments.2.

MATERIAL AND METHODSFor the experimental analysis, we applied an multivariable method (23-1) with no central nor face point, where the number of cycles (repetitions), and solvent absorption were the main variables. On one hand, for the On-Off control method, we also evaluated the solvent absorption target and drying threshold . This study's responses were roughness, geometrical distortion and total spent time. The experiment Design is presented in Table 1, where the main control factors and influence levels were presented.1983All the experiments was done in triplicate in order to statistically supported and the analysis of variance can indicate the confidence level for each variable effect.Table 1 – Design of experimentOn example of total controlled process monitoring is exposed in Figure 2.

In this figure, the target mass absorption, drying threshold and cycles are expressed in function of time.Figure 2 – Schematic of Solvent absorption, minimum threshold, target absorption and number of cycle repetitionFor that process, we have also established a constant evaporation temperature (70°C), while acetone was used as solvent. In this case, the boiling temperature of acetone is 56°C so that we ensure that the vapour generation is kept constant along the exposure time. In addition, the general temperature ramp is presented in Figure 3.Figure 3 – Increasing temperature ramp curveWe have also established that the fabrication parameters were also constants, where: nozzle diameter is 0.4mm, layer height is 0.15mm, object density is 90% and 3 lines of contour shell. The material used was natural ABS GP35For the analysis of surface roughness, we used an optical microscope and image processing in order to obtain the roughness of surface profile.

The material used was natural ABS GP35For the analysis of surface roughness, we used an optical microscope and image processing in order to obtain the roughness of surface profile. The image processing was performed in MATLAB, while the image acquisition have been done by the optic microscope Digital Avangard Optics AN-1984E500 [9], which provides until 500x of amplification magnitude. The geometrical distortion was analysed by a image processing through MATLAB, where the total deviation and local distortions were identified in a resolution of 0.001mm.With respect to the equipment measurement system, the analogue load cell provides an error of 0.001g, while the data acquisition was defined to be 100 samples/s and 12bits resolution .In order to ensure the experiment repeatability we established a procedure that consists in 5 steps, exposure, drying, , stabilization, dimensional analysis, surface roughness analysis, as it is possible to see in Figure 4.

In this figure, the equipment working flow chart is also presented, indicating the operational dynamics of applied ON-OFF control.Figure 4 – Experiment flow chartIn this procedure, we included a stabilisation step before performing the tests after smoothing. This step was included to dry the solvent which impregnated the part surface. Otherwise, the properties of specimens would be jeopardised.It is also important to note that all the specimens that suffered vapour treatment were submitted to microscopic analysis in addition to one specimens which were not exposed to vapour. In this case, these specimens were used as reference to analyze the effects of vapour treatment on the object.3. RESULTS AND DISCUSSIONS With respect to the gravimetric control, we could see that absorption and drying stages imply on cyclic behaviour.

RESULTS AND DISCUSSIONS With respect to the gravimetric control, we could see that absorption and drying stages imply on cyclic behaviour . In this case, further studies are still needed to identify the effect of absorption rates in chemical structure in addition with the correlation with penetration depth and filaments fusing.In the roughness analysis, we identified that the proposed controlled process works in accordance with the expected. Figure 5 shows that the roughness profile reduces the maximum and minimum peaks in addition to reducing Rz and Ra coefficients.1985Figure 5 – Roughness profile analysis in µm dimensionsWith regards to the main effects of control factors on roughness, Figure 6 presents that mass absorption target causes more effect on roughness.

On the other hand, the combination of number of cycles and mass absorption target results in the minimal studied roughness.Figure 6 – Main and secondary effects of control factors on roughnessWe can see that the geometrical distortion was also low in all studied cases, as presented in Figure 7. In this case, the equipment setup that caused minor effect on distortion was below 10μm. It might be explained by the fact that short vapour attack penetrates only superficially and does not significantly alter internal structure of object, as presented in previous studies [1].On the other hand, the polishing appearance also increase as a function of number of passes, as it is possible to be seen in Figure 71986Figure 7 – Polishing effect as a function of number of repetition cyclesIn spite of this study did not evidenced object damages further studies are still needed to identify the process limits.

On the other hand, it was noted that this automatic method reduced human interference and reduces risks of objects damages.4. CONCLUSIONS In this work, the proposed method was identified to work, in addition not to producing severe geometrical distortions nor internal disruption, as hand craft methods.It was possible to determine that this method might reduce the surface roughness in 5 times, besides the geometrical distortion does not exceed 1% even after 3 cycles. We could also identify that higher number of cycle implies on better finishing and automatic mass control tends to avoid potential part jeopardizing, as such long exposure time does.By the end, we can indicate that this sort of process might be included in domestic or small business manufacturing cells, increasing the flexibility, productivity and quality of parts.

This might push industry strategy to change in a short term future and migrate from a centre manufacturing to a distributed manufacturing system.Although this process concept is shown to work, there are still further studies to be done in order to better understand the impact of different smoothing techniques and their control parameter on object properties.5. ACKNOWLEDGMENTSThe authors would like to thank the CNPQ for financial support, as well as the Department of Mechanical Engineering of the University of São Paulo,(campus São Carlos) and the Research department of Concep3D, for providing access to infrastructure and Laboratories.6. REFERENCES1.Cunico, M.W.M., et al., Investigation of additive manufacturing surface smoothing process. Rapid Prototyping Journal, 2017. 23(1): p. 201-208.19872.3.4.5.6. 7.8.9.Priedeman, W.R. and D.T. Smith, Smoothing method for layered deposition modeling, 2003, Stratasys, Inc.: USA. Percoco, G., F. Lavecchia, and L.M.

velopment of Direct SLS Processing for Production of Cermet Composite Turbine Sealing Components· Part n T. Fuestingt, L. Brownt, S. Das*, N. Harlan*, G. Lee*, J. J. Beaman*, D. L. Bourell*, J. W. Barlow*, K. Sargent'RESULTS AND DISCUSSION Indirect SLSCoupons from each of the compositions shown in Table 3 were fired in vacuum to bum off the binder and fully consolidate the coupons. Shrinkage measurements taken from these coupons placed the shrinkage at 4%-6%. Metallographic evaluations revealed that all samples possessed relatively uniform grit distribution. A typical microstructure is provided in Figure 6.Figure 6•. Typical microstructure of indirect processedmaterlai, magnifications lOOX &: 2.50X However Compositions # 4 and #5 illustrated substantial pullout of grit particles during polishing.The microstructures of Composition #4 is provided in Figure 7.

In Composition #4 the grit pull out was associated with the reduced braze content and in Composition #5 this was associated with the alumina used which did not have a titanium coating. This experiment effectively illustrated that the titanium coating on the alumina is necessary. The minimum braze content for an indirect process was established at approximately 18% by weight.Samples of Compositions #2 and # 3 were manufactured according to the process detailedt Allison Engine Company, Indianapolis, Indiana :j: Laboratory for Freeform Fabrication, University of Texas at Austin § Wright Laboratory, Wright-Patterson AFB, Ohio47above. The samples of Composition #3 had circular patches removed which were surface ground and brazed using a braze alloy into the gage section of a single crystal stress rupture bar. The acceptance criteria requires that the sample survive a minimum of 2100'F/35 hours. The results of the four specimens are provided in Table 4.

The acceptance criteria requires that the sample survive a minimum of 2100'F/35 hours. The results of the four specimens are provided in Table 4. Only one sample did not meet the qualification criteria, this sample failed at the braze joint, indicating a bond defect.Figure 7. MicrOStI"llCt1l1-e offnclifeCi Processed Composition # 4 IDustratinggrlt pull-out, magnification 100 XSample3-AM3-BM3-CM3-DMStress (psi) 500500500500Time (lars) 50505050Temperature rF) 2100 48 hrs. 2100 No failure 2100/8.2 hrs. Braze Failure 2100 No failureTable 4 Stress rupture test results of indirect SLS processed cermetroom Rub coupons were also manufactured from the samples and rub tested at temperature. The results are provided in Table 5.

The results are provided in Table 5. A schematic of a rub test is provided in Figure 8.48ProcessingIndirect Indirect Allison conventionalComposition3 2 3 EQuivalentVolumetric Rub Ratio 170 511 140Linear Rub Ratio 13 14 144Table 5 Rub testing of indirect processed material against seal material (Quantabrad 2)Evaluation of room temperature rub testing results on Compositions #2 and #3 provided interesting results. The evaluation involved linear rub ratios which is the ratio of ceramic rub scar depth to abrasive cermet height loss which is not an extremely precise measurement due to the fact that tip height loss, if not significant, is not easily and accurately measured. In general, linear rub ratios in excess of 10 are good and meet abrasive blade tip goals. The indirect SLS materials evaluated met this goal.CMSX-3ABT 3laterialL~~_~~Weiliht Change/DensityMaterial Removed During TesL~ A"••, ' " X ';dlh of RobAbradablc Seal MaterialFigure 8.

illustration of Rub TestingA more accurate measure is a volumetric comparison which is the ratio of rub scar arc depth muhiplied by the rub scar width to the tip weight loss divided by the abrasive density. Simply stated this is the ratio of the abradable seal material rub scar volume to the volume lost on the abrasive tip.As the data indicates, the volumetric comparison provides different results than the linear rub ratio comparison. A volumetric rub ratio in excess of 30 is considered good. The volumetric comparisons indicate the results of the SLS processed material are very good. It is to be noted that all of the above materials would be considered equivalent based on the limited number of tests run.

It is to be noted that all of the above materials would be considered equivalent based on the limited number of tests run. To distinguish one as being better than another, a statistically significant number of lots of materials would need to be conducted and then reviewed with tests from different49accompanying stress rupture data.it was comparable in microstructure and mechanical properties to the material processed by the standard technique. Therefore a final process demonstration was completed to illustrate that production of blade tips via the indirect process is feasible.The testing of indirect SLS processed cermetindicated thatA small lot of material was manufactured to produce 4 pieces of hardware. The parts were laser scanned to final shaPe accounting for 4% shrinkage. The parts were subsequently fired in vacuum. Two of the four pieces of hardware produced are illustrated in Figure 9.

The parts were laser scanned to final shaPe accounting for 4% shrinkage. The parts were subsequently fired in vacuum. Two of the four pieces of hardware produced are illustrated in Figure 9. The microstructure of the parts, presented in Figure 10, was very uniform and passed all requirements. As discussed earlier, follow-on development was aimed at volatilizing the binder using low power scans of the laser. These experiments were canceled due to the coating of laser optics and chamber apparatus by the volatilization of the binder.Figure 10. Typicalllllcrostrllct1lI'e o:f 1D.di.rect Processed blade Tip, Minimal Porosity and Agglomerates, magnification 25X505152SampleK442L240C274N444Stress (psi) 500500500500Time (brs) 50505050Temperature ('F) 2100 No Failure 2100/15 hrs Braze failure 2100 No Failure 2100 No FailureTable 6.

Stress rupture test results of direct SLS processed cennetRub coupons were also manufactured from the samples and rub tested attemperature. The results are provided in Table 7.Processing7/4-A 7/4-B Allison ConventionalCompositionSLS Brazeless SLS Brazeless BrazelessVolumetric Rub Ratio 161 168 187Linear Rub Ratio 31 23 37Table 7. Rub testing of direct processed cemlet against seal material (Quantabrad 2)These results indicate that the brazeless cennet material composition processed by direct SLS is equivalent to the standard cemlet. For sample L240, the failure occurred at the braze joint, in perfomlance attained by elimination of the braze constituent.indicating a bond defect. Additionaltesting is required to establish improvementCONCLUSIONSDirect SLS of cennet abrasive composites and monolithic metallic materials have been developed under this program.

Additionaltesting is required to establish improvementCONCLUSIONSDirect SLS of cennet abrasive composites and monolithic metallic materials have been developed under this program. The processing parameters developed can be modified such that SLS of a variety of fully dense monolithic metallic materials can be conducted. The parameters for SLS of monolithic materials are adjustable such that the resulting microstructure can be tailored to a specific need. Energy densities of approximately 2000 to 4000 J/cm2 are required to eliminate porosity retained around grit particles in the superalloy cemlet composite. For this particular composite material the energy density should be such that the microstructure produced is just barely transfomled to fully dendritic. Energy densities in the range 2500-3500 J/cm2 will produce a fully dendritic microstructure.

Energy densities in the range 2500-3500 J/cm2 will produce a fully dendritic microstructure. Energy densities of 1900 to 2200 J/cm2 will produce a very fine grained equiaxed fully dense superalloy microstructure which may exhibit superplastic properties at elevated temperatures.The mechanical testing results indicate that direct SLS can produce microstructure andproperties equivalent to or better than the conventional labor intensive process.The production of a small lot of abrasive cemlet components and the recycling of all fall off material have proven that the process is production ready and dramatically reduces production costs as outlined in the cost analysis.53roomCompared to the conventional processing method, direct SLS offers the following benefits: 1. Reduced labor and variability by eliminating tape casting and rolling. 2. Repeatable production process. 3. Binderless process. 4. Elimination of lower melting point braze constituent from the composition. 5.

2. Repeatable production process. 3. Binderless process. 4. Elimination of lower melting point braze constituent from the composition. 5. Recyclable fall off material, minimal scrap from any lot size.Indirect SLS offers the following benefits: 1. Reduced labor and variability by eliminating tape casting and rolling. 2. Repeatable production process. 3. Reduced labor in obtaining the correct thickness and porosity. 4. Ability to manufacture much smaller lot size as compared to conventional manufacturingmethod.The indirect SLS process has some disadvantages as compared to the direct SLS process. These are: 1. Fall off material may possibly not be recyclable. 2. Requires handling the parts in the delicate green state which may produce additional scrap.The next stage of research under this program will focus on developing a high Yield,rePeatable production process.ACKNOWLEDGEMENTSPartial funding for this work was provided by the United States Air Force, Contract No.

F33615-94-C-2424. The remainder of the resources were provided by Allison Engine Company and the University of Texas at Austin.REFERENCES1 SIdra, Charles A., Cost Reduction ofAdvanced Turbine Engines, 31stAIAAlASMFJSAEIASEE Joint Propulsion Conference and Exhibit, July 1995.2 Proceedings of the First Air Force, Wright Laboratory S&T Affordability Exit Criteria Workshop.3 Dix, D.M. and.Riddell, FR., Projecting Cost-Performance Trade-of/sfor Military Vehicles, Attachment to Anita Jones letter to JAST program office on IHPIET AFFORDABILTI, dated 31 August 1994, Aeronautics & Astronautics, September 1976.4 Yeaple, F., Gas Turbine Rotor Grinds Own Tip Seals, Design News, 1-5-87, pp. 106-107.5 Helms, Harold E., Heitman, Peter W., Lindgren, Leonard E. and Thrasher, Samuel R., Ceramic Applications in Turbine Engines, Noyes Publications, Park Ridge, New Jersey, 1986, pp. 131-137.6 Beaman, Joseph J.

VELOPMENT OF EXTRUSION-ON-DEMAND FOR CERAMIC FREEZE- FORM EXTRUSION FABRICATIONThomas Oakes, Parimal Kulkarni, Robert G. Landers, and Ming C. LeuDepartment of Mechanical and Aerospace EngineeringMissouri University of Science and Technology,Rolla, MO 65409{tmo6w3@mst.edu, pskf44@mst.edu, landersr@mst.edu, mleu@mst.edu }{Oakes: (573) 341-4862, Kulkarni: (573) 263-5949, Landers: (573) 341-4586, Leu: (573) 341-4482}ABSTRACTIn the Freeze-form Extrusion Fabrication (FEF) process, extrusion-on-demand (EOD) refers to the ability to control the start and stop of paste extrusion on demand and is vital to the fabrication of parts with complex geometries. This paper describes the development of EOD for ceramic FEF through modeling and control of extrusion force, selection of appropriate process parameters, and a dwell technique for start and stop of extrusion.

A general tracking controller with integral action is used to allow tracking of a variety of reference forces while accounting for the variability in the paste properties. Experiments are conducted to model the process and tune the controller. The developed technique for EOD is demonstrated to fabricate a number of cross sections and three- dimensional parts from alumina paste.1. INTRODUCTION Fabricating ceramic materials into usable 3D components is a complicated, costly, and time-consuming process that usually involves pressing ceramic powder to near net shape, sintering, and finish machining or diamond grinding of the final geometry from a block of dense ceramic. Using organic or aqueous based slurries, cast into a prefabricated mold designed to mimic the shape of the final product, is another method for fabricating 3D ceramic components. After casting, the green form undergoes binder removal and sintering to produce a dense ceramic product.

After casting, the green form undergoes binder removal and sintering to produce a dense ceramic product. The cost and time associated with mold design and processing tend to increase greatly with geometric complexity of the component. Further, a mold-based fabrication process cannot produce 3D components with internal passages and cavities. In recent years, several Solid Freeform Fabrication (SFF) techniques have been developed to fabricate 3D ceramic components, including Fused Deposition of Ceramics [1], Fused Deposition Modeling [2], Extrusion Freeform Fabrication [3], 3D Printing [4], Selective Laser Sintering [5,6], Selective Laser Melting [5], Shape Deposition Manufacturing [7], and Robocasting [8,9]. All of these techniques involve layer-by-layer material addition. Most of them produce a nearly finished product (after sintering) without a mold, while some others involve fabricating a mold or 3D tooling to form the final component by casting.

Most of them produce a nearly finished product (after sintering) without a mold, while some others involve fabricating a mold or 3D tooling to form the final component by casting. Especially relevant to the research in the present paper are the Extrusion FreeformFabrication (EFF) and Robocasting processes. EFF is the first SFF process in utilizing extrusion of organic pastes to build 3D ceramic components [3]. Robocasting is the first SFF process that extruded aqueous pastes to fabricate parts from ceramics [8,9]. The use of aqueous based extrusion has several unique advantages: (1) it is environmentally friendly; (2) its minimal amount of organic binder reduces time for binder pyrolysis; (3) its equipment cost is low compared with laser based and other SFF processes; and (4) the process can easily accommodate use of multiple materials to make functionally graded components.EFF and Robocasting are room temperature SFF processes.

Due to the intrinsicproperties of ceramic pastes in room temperatures, deformation (slumping) of large parts during the fabrication process cannot be avoided. One way to remedy this is to solidify the paste rapidly in a freezing environment after it is extruded. This can be accomplished by extruding aqueous pastes at temperatures below the freezing point of the aqueous medium. This is the basis of the Freeze-form Extrusion Fabrication (FEF) process, which extrudes aqueous ceramic paste of high solids loading in a layer-by-layer manner for part fabrication. The FEF process has been demonstrated for fabrication of ceramic parts by continuous extrusion of aqueous paste [10,11].The FEF study described in the present paper uses ceramic paste made of alumina(Al2O3) powder, water, dispersant, lubricant and other materials. The ceramic solids loading is up to 50 vol.% of the paste. Water is the main liquid medium, with the organic binder content only 2–4 vol.%.

The ceramic solids loading is up to 50 vol.% of the paste. Water is the main liquid medium, with the organic binder content only 2–4 vol.%. A ram extruder mechanism is used to extrude the paste. The extruder is mounted on a 3D gantry system. The experimental setup is shown in Figure 1. The green part obtained after fabrication is freeze-dried, the binder is removed through a burnout process, and the brown part is then sintered to obtain the final part.Figure 1: Experimental setup with an extruder ram mounted on a 3D gantry systemControl of the paste flow rate, extrudate consistency, and start and stop of extrusion as prescribed is a major challenge in the FEF process development. This requires understanding of the paste dynamics. Benbow and Bridgewater [12] developed an equation relating the extrusion pressure to the paste velocity during steady-state extrusion, with the assumption that the paste is incompressible and undergoes slip in the die land. Mason et al.

Mason et al. [13] conducted an empirical study on the dynamics of the paste and showed that the paste flow can be modeled as a first-order system with significant variations in the gain and time constants. Amarasinghe and Wilson [14] stated that ceramics are generally much more difficult to extrude as compared to polymers due to temporal and spatial inconsistencies in the paste properties. Another study investigated the use of a bang-bang control strategy with feedback control of the extrusion force [15]. Drawbacks of this control technique include excessive use of trial and error and lack of systematic protocols. Zhao et al. [16] successfully conducted continuous extrusion fabrication with an adaptive controller to regulate the extrusion force. However, the resulting parts were limited to shapes that could be made with continuous extrusion. Another study by Mason et al. [17] investigated start and stop of extrusion in conjunction with force control.

Another study by Mason et al. [17] investigated start and stop of extrusion in conjunction with force control. This paved the initial path for the extrusion-on-demand extrusion study presented in the present paper.The objective of this paper is to develop a force controller that is capable ofextrusion-on-demand for the FEF process. The rest of the paper is organized as follows. In Section 2, the FEF experimental setup and process modeling are described. Next, in Section 3, the dynamic paste extrusion model is used to design a general tracking controller with integral action to account for paste variability. The main FEF process parameters are introduced in Section 4, along with the acceptable ranges of these parameters determined experimentally. The fabrication of 2D cross-sections and 3D parts with the developed extrusion-on-demand FEF process is described in Section 5. Finally, a conclusion of the study is given in Section 6.2.

The fabrication of 2D cross-sections and 3D parts with the developed extrusion-on-demand FEF process is described in Section 5. Finally, a conclusion of the study is given in Section 6.2. PROCESS MODELING A series of tests are conducted to obtain a dynamic model for the FEF extrusion process. The input is set to an upper bound voltage (250 mV) until the start of extrusion has occurred as is visually confirmed. Once extruding, the system is subjected to a series of step changes in the command voltage which is cycled several times. Plots of the response and input are shown in Figure 2. (1)2. PROCESS MODELING A series of tests are conducted to obtain a dynamic model for the FEF extrusion process. The input is set to an upper bound voltage (250 mV) until the start of extrusion has occurred as is visually confirmed. Once extruding, the system is subjected to a series of step changes in the command voltage which is cycled several times.

Once extruding, the system is subjected to a series of step changes in the command voltage which is cycled several times. Plots of the response and input are shown in Figure 2. =where b and a are model coefficients to be determined through a recursive least squares algorithm. Equation (1) is transformed into the difference equation( bU i The recursive least squares (RLS) algorithm is implemented with ]T ) 1( ) F i( aF i) 1 − +) − 1= −[ −=ϕ( ) ( 1 iU iF − [ ]Tba 0=η−0where a0 = –0.998 and b0 = 0.0397, which are determined empirically based on the RLS algorithm. Equation (1) is thus .0 0397 z 998.0 −K b ) 1( − exp( T −( ) zG==z−)/ τwhere K = 0.0413 kN/mV, T = 0.1 s is the sample period, and τ = 42.7 s is the time constant. The model is used to plot against the data collected from the experiment as shown in Figure 2.

The model is used to plot against the data collected from the experiment as shown in Figure 2. The good agreement indicates that the first-order process approximation is acceptable to describe the process dynamics and can be used for controller design.1000)Nmeasurementmodel( e c r o f5000 0100200300400500600700800200)N(r o r r e0200 0100200300400500600700800)V m100( e g a t l o v0100 0100200300400 time (s)500600700800Figure 2: Comparison of measurement data with modeled response for the given inputprofileTo find the relationship between ram extrusion force and extrudate velocity at steady state, a series of tests are conducted to collect values of steady-state extrudate(2)(3)(4)(5)velocity at constant ram extrusion forces. Based on the experimental data, the force- velocity relationship is approximately linear over the range of forces used in the experiment.

Based on the experimental data, the force- velocity relationship is approximately linear over the range of forces used in the experiment. A least squares regression results in the following relationship:Vext=.0043F−.7403where Vext is the extrudate velocity in m/s. The regression yields R2 = 0.857, indicating an acceptable correlation.3. CONTROLLER DESIGN A control algorithm is designed to allow for extrusion on demand of the FEF process in order to extrude the paste at a constant rate and to coordinate the start of extrusion with the gantry motion. This algorithm uses the general tracking control to reject constant disturbances with desired error dynamics − 1−= z3. CONTROLLER DESIGN A control algorithm is designed to allow for extrusion on demand of the FEF process in order to extrude the paste at a constant rate and to coordinate the start of extrusion with the gantry motion.

This algorithm uses the general tracking control to reject constant disturbances with desired error dynamics − (8)( ) zE( ) zR( )zFwhere R(z) is the reference ram extrusion force and F(z) is the measurement from the load cell. a(z) is the denominator of the open-loop transfer function, and g(z) is a first- order polynomial ( ) zg=zg 1+g0where g1 and g0 are determined by the desired closed-loop error dynamics. Equation (7) can be rewritten as ( ) a zg + 1[ z] ( ) 0 ) zEg =( −−+21 −−a0In order to allow the system to exhibit a first-order response, one of the poles in the control system is suppressed by making the second pole at least one order of magnitude smaller than the dominant pole. The dominant pole is selected in order to decrease the settling time as much as possible without causing system instability. Two overdamped poles are thus selected with time constants τ1 = 1.5 s and τ2 = 0.15 s.

The controller tracks a step reference with an approximate settling time of 18 s. This response time is much too slow for extrusion on demand so the controller is implemented again with the reference force ramped to decrease the settling time. The result is shown in Figure 3, which demonstrates that the controller tracks the ramp reference with an approximate settling time of 5 s. Thus using a ramp reference force rather than a step reference force has decreased the response time by a factor of more than 3. It should be noted that the controller does not(6)(9)(10)(11)(13)(14)saturate throughout the test due to smaller changes in the error. The pulsing of the ram caused by the constant over and under correction during the ramp input probably excites the paste into a faster response.

The pulsing of the ram caused by the constant over and under correction during the ramp input probably excites the paste into a faster response. From this point forward, all reference forces are set to ramp to the desired values with the same slope.)N200responsereference( e c r o f1000 02468101250)N(r o r r e050 024681012)V m500(lo r t n o c0500 0246 time (s)81012Figure 3: Response of ram extrusion force for a ramp inputThe controller is subjected to a variety of changes in the reference force to compare their effects on variation in the control signal and, as a result, the ram velocity required to maintain different reference forces. It has been experimentally established that reference forces greater than 535 N will result in controller instability.4.

It has been experimentally established that reference forces greater than 535 N will result in controller instability.4. PROCESS PARAMETERS The quality of the FEF process is affected by many process parameters including the table velocity, ram velocity, dwell periods, standoff distance, overlap factor, and environmental and heating temperatures. This section discusses these parameters and a series of experiments conducted to determine the ranges of working values for these parameters for the extrusion-on-demand FEF process. The nozzle diameter for all experiments is 580 microns. 4.1. Ram Velocity The ram velocity is controlled directly by the command input voltage and is constantly updated online to regulate the extrusion force. The ram velocity and the command input voltage are positively correlated, implying that the ram velocity and the extrusion force are also positively correlated.

The ram velocity and the command input voltage are positively correlated, implying that the ram velocity and the extrusion force are also positively correlated. In general, a constant ram velocity produces a fairly constant force, but only if the paste has already begun to extrude out of the nozzle. The ram velocity continuously oscillates to maintain a constant force. In general, the average ram velocity over a period of steady-state force control has a slightly positive bias, generally in the order of μm/s.4.2. Table Velocity The magnitude of the table velocity should be set equal to the magnitude of the extrudate velocity, which is the velocity of the paste as it exits the nozzle, to ensure that paste is distributed uniformly. Equation (6) empirically relates the average extrudate velocity to the average input ram force. This relationship implies that a constant force input will result in a constant extrudate velocity, and that the two are positively correlated.

This relationship implies that a constant force input will result in a constant extrudate velocity, and that the two are positively correlated. However, a point exists at which there is not enough force to cause extrusion. Table velocities that are too high with respect to the extrudate velocity will result in thin and sometimes incomplete paste deposition, thus forming voids. On the other hand, table velocities that are too low result in excessive deposition per unit length, causing surface defects from dragging the nozzle and forming thick, non-uniform lines.4.3. Standoff distance refers to the distance between the nozzle tip and the top of a previously deposited layer (or substrate). It is expressed as percentage of the measured extrudate diameter. A standoff distance that is too low will result in thick extrusion containing a surface defect due to the nozzle dragging through the previously deposited layers, possibly resulting in part collapse.

A standoff distance that is too low will result in thick extrusion containing a surface defect due to the nozzle dragging through the previously deposited layers, possibly resulting in part collapse. A standoff distance that is too high will result in discontinuous or thin lines during extrusion due to lack of contact between the paste and surface, eventually leading to uncontrolled deposition and part collapse.Standoff DistancePrior to determining a working range of standoff distance, the extrudate diameter must be determined. This is accomplished by photographing the extrudate during steady- state extrusion and pixel comparison of the photograph. The experiment is conducted over constant ram extrusion forces ranging between 350 N to 500 N in increments of 25 N.

The experiment is conducted over constant ram extrusion forces ranging between 350 N to 500 N in increments of 25 N. The average extrudate diameter is calculated to be 617 μm with a standard deviation of 6.44x10-2 μm.Single line testing is not reliable in this case because inappropriate standoffdistances are generally not made apparent until several layers have been deposited. A series of four hollow cylinders 25.4 mm in height and 25.4 mm in diameter are fabricated at the ram extrusion force of 450 N and the table velocity of 4.87 mm/s. The tested standoff distances are 70%, 55%, 40%, and 25% of the extrudate diameter, and the result is shown in Figure 4. The standoff distance at 70% is too high for good layer consistency. The parts fabricated at 55% and 40% both appear to have good layer consistency and surface finish. The cylinder fabricated at the standoff distance of 25% of extrudate diameter contains no air gaps, but is visually much thicker than the rest of the specimens.

The cylinder fabricated at the standoff distance of 25% of extrudate diameter contains no air gaps, but is visually much thicker than the rest of the specimens. The established range of standoff distances is thus between 40% and 55% of the extrudate diameter.Figure 4: Testing of four standoff distances: 70%, 55%, 40%, and 25% of the extrudatediameter4.4. Overlap Factor The overlap factor refers to the spacing between two neighboring lines and is defined asKOL=−dgapDextwhere Dext is the extrudate diameter and dgap is the gap between two neighboring deposited lines. An overlap factor of zero implies that the centers of two successive parallel lines are exactly one extrudate diameter apart. A positive overlap factor implies that the line centers are closer than the extrudate diameter, which can lead to accumulation of material and undesirable part quality. A negative overlap factor implies that the line centers are further apart than one extrudate diameter.

A negative overlap factor implies that the line centers are further apart than one extrudate diameter. Parts with excessive negative overlap factors will have voids between extruded lines, which may lead to poor structural integrity and will ultimately lead to the part collapse.A series of single-layered plate, 25.4 mm x 35.6 mm, are fabricated at a standoff distance of 55% of the extrudate diameter to determine the working range of the overlap factor. All tested values of overlap factor are negative due to the compaction of deposited material caused by the low standoff distance (55% of the extrudate diameter). Values tested are -60%, -45%, -30%, and -15% with three cross-sections fabricated per overlap factor. The result is shown in Figure 5. Rastered parts at -60% contain visual voids in the rastering structure. Parts created at -15% contains slight accumulation of materials between deposited lines, implying the spacing is too close.

Rastered parts at -60% contain visual voids in the rastering structure. Parts created at -15% contains slight accumulation of materials between deposited lines, implying the spacing is too close. Overlap factors of -45% and - 30% both exhibit acceptable raster quality.Figure 5: Testing of four (4) overlap factors: -60%, -45%, -30%, and -15% of theextrudate diameter; standoff distance = 55% of the extrudate diameter4.5. Dwell Time for Start and Stop Determining the dwell time at the start and stop of extrusion is important for extrusion- on-demand (EOD). Dwell time refers to the amount of time required for the gantry to remain stationary after the force controller is activated for paste extrusion. Excessive dwell times at the start cause material accumulation at the beginning of the line and can lead to nozzle clogging.

Excessive dwell times at the start cause material accumulation at the beginning of the line and can lead to nozzle clogging. Dwell times that are too short at the start may cause the deposition head to begin advancing along the path prior to extrusion of paste, leading to incomplete deposition lines.Start dwell time is set as a function of the reference ram force. The gantry isallowed to advance along its path after a specified force (in percentage of the reference(15)ram force) is attained by the force controller. Stop dwell time is simply set to three seconds, in which the gantry movement is terminated followed by the ram retraction at the maximum motor speed prior to activation of the next motion command.Since the start dwell time is a function of the extrusion force, it is also a function of the extrudate velocity (which is also the table velocity).

Good EOD is hard to achieve at high speeds, and the performance of the start and stop technique improves as the velocity/force is decreased. However, operating at a very low velocity/force combination is not recommended due to potential for discontinuous extrusion and nozzle clogging besides low productivity.The start dwell time for the FEF process is analyzed through a series of extrusion line tests. The reference extrusion force is varied from 350 to 500 N in 25 N increments at a standoff distance equal to 55% of the extrudate diameter for each test. During each test, five 127 mm lines are extruded at dwell times ranging from 50% to 70% in 5% increments of the reference ram force. The result is shown in Figure 6, which depicts that the amount of excess material accumulation during the dwell period decreases as the reference extrusion force decreases.

The result is shown in Figure 6, which depicts that the amount of excess material accumulation during the dwell period decreases as the reference extrusion force decreases. This implies that a low ram force should be used for extrusion; however, a low force does not guarantee continuous extrusion over time. In most cases, lines deposited with dwell times set at less than 65% of the extrusion force are discontinuous. All lines deposited with dwell times set at larger than 70% of the extrusion force display moderate to severe tapering. The start dwell time is thus established between 65% and 70% of the reference extrusion force.Figure 6: Lines deposited with the start dwell time ranging from 50% to 70% in 5%increments of the reference extrusion force4.6. Environmental and Paste Temperatures The environmental temperature refers to the temperature of the freezer in which the gantry system is contained.

Environmental and Paste Temperatures The environmental temperature refers to the temperature of the freezer in which the gantry system is contained. Liquid nitrogen is dispersed in streams aimed at the substrate, keeping the environment temperature at -5 °C around the build area. As the environmental temperature decreases, the process allows for faster freezing of deposited material and higher part stability. Temperatures that are too low may result in nozzle clogging due to freezing of paste in the die land. Increased environmental temperatures allow for higher extrudate velocities at lower forces, but reduce part stability due to insufficient freezing of deposited lines.The paste temperature refers to the temperature of the paste maintained by theheating tape around the sleeve of the stainless steel syringe. Low paste temperatures may result in partial freezing of the paste inside the syringe, preventing extrusion at lowerforces.

Low paste temperatures may result in partial freezing of the paste inside the syringe, preventing extrusion at lowerforces. High paste temperatures change the paste consistency, making it more difficult to quick freeze the extrudate. The paste temperature is fixed at 10 °C in this study.5. PART FABRICATION Experiments are conducted to fabricate 2D cross-sections and 3D parts by the FEF process requiring EOD. The process parameters used for the fabrication are: table velocity = 4.87 mm/s, standoff distance = 55% of the extrudate diameter, dwell time = time arriving at 70% of the reference ram extrusion force, and overlap factor = -45% of the extrudate diameter. Insight 4.3.1 software is used to slice .stl files into sections and generate extrusionpaths based on the selected process parameters. The layer and generated tool path of a simplified fuel injector strut is shown in Figure 7.

The layer and generated tool path of a simplified fuel injector strut is shown in Figure 7. The part is scaled to 7.62 mm x 50.8 mm x 25.4 mm and oriented such that the larger dimensions are on the x-y plane, in order to minimize the build height.Figure 7: Generation of layers and tool paths using InsightThe fuel injector strut’s cross sections are deposited in a side-by-side manner firstto test the viability of continuous extrusion between commanded deposition starts and stops. Each layer contains two starts and stops, the first for contouring and the second for internal rastering. The results are shown in Figure 8. The experimental setup is capable of continuous extrusion over all layers. Material accumulation is observed at all points of initial deposition, implying that stacking of these layers could result in a height offset which would interfere with the standoff distance.

Material accumulation is observed at all points of initial deposition, implying that stacking of these layers could result in a height offset which would interfere with the standoff distance. The second cross-section down from the top right contains a void due to an air gap in the syringe.Layers of alumina paste are deposited and stacked to fabricate a simplified fuelinjector strut. Successive layers are oriented 90° apart to improve the structural stability. Figure 9 shows four of these parts fabricated by the process.Figure 8: Extrusion of 15 cross-sections of a simplified fuel injector strutFigure 9: Simplified fuel injector struts fabricated by the FEF process6. CONCLUSIONS Working ranges have been established for FEF of aqueous alumina paste for the following process parameters: table velocity, ram velocity, standoff distance, dwell time, overlap factor, environmental temperature, and paste temperature.

Extrusion on demand has been made possible through the implementation of a general tracking controller with appropriate start and stop dwell times at the beginning and end of a continuous extrusion path. Successful 2D cross sections and 3D parts have been fabricated using the FEF process with the developed EOD control technique. Further study of the extrudate velocity profile along the extrusion path will also be needed in order to eliminate excess accumulation of material at the beginning of extrusion.7. ACKNOWLEDGMENTS This research was supported by the U.S.

ACKNOWLEDGMENTS This research was supported by the U.S. Air Force Research Laboratory through the Center for Aerospace Manufacturing Technologies (CAMT), contract #FA8650-04-C- 5704, at the Missouri University of Science and Technology (Missouri S&T), and byCAMT Industrial Consortium whose members consist of Boeing, General Dynamics, Spirit AeroSystems, GKN Aerospace, Siemens, KMT Waterjet, the Product Innovation and Engineering, Spartan Light Metal Products, and the Third Wave Systems. The research was also supported by Missouri S&T’s Intelligent Systems Center. Special thanks are due to Drs. G. E. Hilmas and T. Huang for expert advice on alumina paste preparation and post processing.8. REFERENCES [1] Bandyopadhyay, A., Panda, P. K., Agarwala, M.K., Danforth, S.C., and Safari, A., “Processing of Piezocomposites by Fused Deposition Technique,” Journal of the American Ceramic Society, Vol. 80, No.6, pp. 1366-1372 (2000). Patent No.

velopment of Freeform Master I - a desktop RP machine based on a new sheet lamination processKwan H. Lee and Joung O. ParkDepartment of Mechatronics Kwangju Institute of Science and Technology (K-JIST) Kwangju, KoreaAbstractA novel process was developed for building Rapid Prototyping(RP) parts using a sheet lamination technique. The building process of existing sheet lamination RP machines consist of the following steps : feeding, lamination and cutting. In this process, the laminated part of an object is often scratched by a cutter or damaged by a laser beam due to the cutting operation after the lamination step. In addition, decubing of the unused portion from the laminated block is difficult. In the new process, however, cutting is performed before lamination. The cutting operation takes place while a paper sheet is firmly attached, using electrostatic force on the plate. Then liquid glue is applied only to the calculated region of the given contour for lamination.

Then liquid glue is applied only to the calculated region of the given contour for lamination. This new process aims to manufacture a $2k RP machine, what we call the Freeform Master I and can use A4 size or letter size sheets of paper. A prototyping machine that demonstrates the concept design was built and further research issues are discussedIntroductionThe rapid prototyping(RP) process is a well known technique that build materials in layers and manufactures some designed form with the cross-sectional data in three-dimensions. This kind of equipment based on RP technique have many merits, such as they can manufacture various forms or shapes. Such machines and the materials they use, however, are very expensive. Therefore, the need for an inexpensive RP machine has increased recently. According to this need, we tried to develop an inexpensive RP machine that uses cheaper materials.

Therefore, the need for an inexpensive RP machine has increased recently. According to this need, we tried to develop an inexpensive RP machine that uses cheaper materials. Therefore the sheet lamination RP process was chosen to this aim.In a general RP process, we get information of the cross sections from the corresponding 3D CAD model, the practical model is made through various RP processes and then the unnecessary parts are removed. This research is focused on the process of making a practical RP model.The current sheet lamination RP processThe existing sheet lamination RP process is discussed in the following steps:2831) Feeding - There are generally two types of paper used in the sheet lamination RP process.

Rolling paper is used in process 1 of Fig.1 and it is fed by a roller, and plain paper is used in process 2, and it is also fed by a roller, 2) Gluing & Lamination – The gluing & lamination methods between the two processes are different in accordance with the method of feeding the paper. In process 1, the heat-melted glue is pasted on one side of the paper and the paper is pressed by a heated roller. In process 2, the glue is printed on a piece of plain paper in a laser printer and the papers are pressed with a heated plate. 3) Cutting - In process 1, a laser beam is used and in process 2, a knife is used for cutting. lamination - cutting.Fig.

1 Two types of sheet lamination processThe problems of the current sheet lamination RP process and solutionsThe problems of the current sheet lamination RP process are summarized as follows:1) Gluing of the unnecessary part - The fact the cutting step is after the gluing & lamination step mean that gluing the unnecessary part, regardless of the need.(process 1) _ Perform the gluing & lamination step after the cutting step.2) The damage of existing layers - The layers which have already been made can be damaged by cutting after the gluing & lamination step.(process 1, process 2) _ Perform cutting before the gluing & lamination step.3) Fixing the first layer of materials and removing them - There are some problems in fixing the first layer of materials and detaching it from the plate since cutting is done after gluing & laminating in the current sheet lamination RP process.(process 1, process 2) _ Fix and detach the first layer of materials by using an electrostatic force.284The requirements of the new sheet lamination RP processTo solve the problems of building the unnecessary part and damaging the existing layers, we changed the working process from a 'cutting step after the gluing & lamination step' to 'the gluing & lamination step after cutting'.

To solve the problem of fixing the first materials we fix and detach the first layer of materials using an electrostatic force. Overall, we changed the existing sheet lamination RP process from 'feeding _ gluing & lamination _ cutting' to 'feeding _ cutting _ gluing & lamination' and defined the basic functions as followings.uuFeeding - feed A4 size paper. Cutting - cut the paper or materials.u Gluing & lamination - glue and press the paper or materials after spraying the glue onlyon the necessary part.Fig. 2 Overview of the proposed processTo develop this new process, we arranged the necessary conditions as follows:1) Feeding. ·Sheets of paper must be fed one after another. ·Transportation to the place of manufacturing must be easy.2) Cutting ·A place for cutting is necessary. ·We must fix the paper so that it can be firmly attached to the plate while cutting.3) Gluing & lamination ·The glue must be sprayed only to the necessary part selectively.

·We must fix the paper so that it can be firmly attached to the plate while cutting.3) Gluing & lamination ·The glue must be sprayed only to the necessary part selectively. ·The materials must be pressed for safety adhesion. ·A plate to fix the first layer is needed.The process and device for the new sheet lamination RPThe detailed process based on the preconditions is as follows:1) Feed the materials - feeding device 2) Carry the materials to the place of cutting and glue spraying - conveying device 3) Cut the materials - cutting device 4) Spray the glue selectively only to the necessary region which is already built - spraying device2855) Build the materials - lamination deviceWe modularized all the devices after we organized all the necessary steps. The cutting device(3) and the spraying selectively device(4) are similar in their functions, because they are operated along the cross sections from a 3D CAD model. So we united the two devices.

The cutting device(3) and the spraying selectively device(4) are similar in their functions, because they are operated along the cross sections from a 3D CAD model. So we united the two devices. Therefore, all the process is operated through 4 modules; feeding device, conveying device, cutting/spraying device and lamination device.(1) Feeding deviceThe feeding device is used to feed the materials into the unit. If the user of this machine feeds materials in preparation of the whole process, then this device will help by feeding the papers one sheet after another to the cutter. This feeding device makes the shaked papers move up and down so that it can carry one sheet of paper after another by the electrostatic panel.(2) Conveying deviceThis device carries papers from the feeding device to the lamination device.

It carrieseach sheet of paper after it has been fixed to the electrostatic plate by the electrostatic force.(3) Cutting/spraying deviceThis device cuts the piece according to the cutting information the user inputs after the paper has been fixed tightly on the plate by the electrostatic force. After cutting, the glue is selectively sprayed on only the necessary region according to the information the user inputs. Cutting is performed with knife, and spraying of the glue is performed through a nozzle using the vapor pressure.(4) Lamination deviceAfter the cutting and spraying step is finished, the lamination device moves up and down, and press the cut materials on the part where the glue have sprayed.

The electrostatic force is used to fix the first layer and laminated papers until the whole building process is finished.A rapid prototyping equipment applying the electrostatic force(Freeform Master I1)The real size of the Freeform Master I which was made based on these functions, is 1070 mm _ 760 mm _ 720 mm and its working space is 247 mm _ 160 mm _ 150 mm. The prototype machine developed is shown in Fig. 3.1 The proposed RP machine based on the new sheet lamination process is named as the Freeform Master I.286Fig. 3 The Freeform Master I(left) and the CAD model of Freeform Master I (right)Fig. 4 The CAD model (left) and manufactured parts(right) by Freeform Master IThe time required to build parts with this machine differs according to the shape and complexity of the part.

4 The CAD model (left) and manufactured parts(right) by Freeform Master IThe time required to build parts with this machine differs according to the shape and complexity of the part. It took 120 seconds to finish one cycle of the process in making simple shaped part as shown in Fig.4.ConclusionIn our prototype machine, we simplified the RP process and reduced the cost and the complexity of the system using an open-loop control. The new process proceeded using a 'lamination step after cutting', while any materials can be used if they are nonconductive materials, because they need to be fixed by an electrostatic force. Moreover, we can use any used papers in this machine, if they are not damaged in their size and shape.

Moreover, we can use any used papers in this machine, if they are not damaged in their size and shape. Therefore, we are sure that this machine can contribute in the reduction of the expense as well as it can contribute for reusing of resources.We are going to further develop this machine to make it more convenient and reliable so that it can be used in daily life and as a teaching aid in schools. Moreover, if the bulk of it is reduced, we can use it as a peripheral, too.References1. Sandig, S., Wiesner, P., 1999, “Manufacturing of metallic prototypes and tools by laser cutting and diffusion bonding”, Intelligent Processing and Manufacturing of Materials, IPMM '99. Proceedings of the Second International Conference on Volume: 1 , Page(s): 251 -253 vol.12. Kyonghwan No, 1998, “Development of 3D FAX using lamination technique” 3. Hakgeun Jeong, 1998, “Development of an electrically suspended transport system using the self-sensing mechanism”4. Kai, C. C.and Fai, L.

VELOPMENT OF FREEZE-FORM EXTRUSION FABRICATION WITH USE OF SACRIFICIAL MATERIALMing C. Leu and Diego A. GarciaDepartment of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractThe development of Freeze-form Extrusion Fabrication (FEF) process to fabricate three- dimensional (3D) ceramic parts with use of sacrificial material to build support sections during the fabrication process is presented in this paper. FEF is an environmentally friendly, additive manufacturing process that builds 3D parts in a freezing environment layer-by-layer by computer controlled extrusion and deposition of aqueous colloidal pastes based on computer-aided design (CAD) models. Methyl cellulose was identified as the support material, and alumina was used as the main material in this study.

Methyl cellulose was identified as the support material, and alumina was used as the main material in this study. After characterizing the dynamics of extruding alumina and methyl cellulose pastes, a general tracking controller was developed and applied to control the extrusion force in depositing both alumina and methyl cellulose pastes. The controller was able to reduce the time constant for the closed-loop system by more than 65% when compared to the open-loop control system. Freeze-drying was used to remove the water content after the part has been built. The support material was then removed in the binder burnout process. Finally, sintering was done to densify the ceramic part. The fabrication of a cube-shaped part with a square hole in each side that requires depositing the sacrificial material during the FEF process was demonstrated.1. IntroductionCeramic materials are applied widely in aerospace, automotive, biological, and other industries [1].

IntroductionCeramic materials are applied widely in aerospace, automotive, biological, and other industries [1]. Many ceramic materials, such as Al2O3 and ZrB2, can survive high temperatures (up to 2000°C for alumina and 3000°C for zirconium diboride), but processing these materials for use as 3D components is often challenging, expensive, and time-consuming. Building a ceramic part using additive manufacturing (AM) may reduce the material cost and build time for small runs and for parts with complex geometries.Several AM processes can be used to produce ceramic parts. One is Fused Deposition of Ceramics (FDC) [2-5], which is based on the Fused Deposition Modeling (FDM) process invented by Scott Crump [6]. Stereolitography (SLA) [7,8], 3D Printing (3DP) [9], and Selective Laser Sintering (SLS) [10,11] are commercialized AM techniques for fabricating mostly polymer components, but with limited capabilities to make ceramic parts.

Most existing paste extrusion based additive manufacturing systems utilize non-aqueous ceramic-based materials, and require a large amount of organizer binder for part fabrication (40-50 vol.%). Robocasting, which is an aqueous based extrusion freeform fabrication process, has been used to produce parts from different types of ceramics including oxides and non-oxides, and biomaterials [12]. However, it operates at room temperature, and has difficulty of preventing deformation (slumping) of large parts during the fabrication process.One way to part deformation during the fabrication process is to freeze the paste rapidly after it is extruded and deposited. This can be accomplished by extruding aqueous-based pastes (with only a small amount of organic binder) at temperatures lower than the freezing point of the aqueous medium.

This can be accomplished by extruding aqueous-based pastes (with only a small amount of organic binder) at temperatures lower than the freezing point of the aqueous medium. Freezing during extrusion is the basis of the Freeze-form Extrusion Fabrication (FEF) process, which is a layer-by-layer deposition process that was developed at Missouri S&T by extending the technology of Rapid Freeze Prototyping (RFP) [13-15]. This environmentally friendly process has been developed and demonstrated for the freeform fabrication of 3D ceramic-based components. The aqueous paste used in the FEF process is extruded by a ram extruder, and the extruded material immediately deposits on a working surface (which may be a substrate or a previously deposited layer) that can be moved by an X-Y table. The surface temperature is set to a sub-zero temperature to freeze the material as it is deposited.

The surface temperature is set to a sub-zero temperature to freeze the material as it is deposited. The solids loading of the ceramic paste could be in the range of 45-55 vol.%.The process parameters required to achieve better part quality in the FEF process have been studied previously [16-17]. Huang et al. [16] designed and implemented an on-off feedback controller to achieve more consistent material extrusion. This controller used the reading of extrusion force from a load cell to automatically adjust the ram velocity. Zhao et al. [17] found that due to effects such as air bubble release, agglomerate breakdown, and liquid phase migration, the ram velocity was difficult to control.

Zhao et al. [17] found that due to effects such as air bubble release, agglomerate breakdown, and liquid phase migration, the ram velocity was difficult to control. Hence, an adaptive controller was designed and implemented to regulate the extrusion force.The objective of the research as described in the present paper was to develop a method of fabricating ceramic parts that require the use of sacrificial material in the FEF process, which has not been investigated in the previous studies. Choosing a suitable support material for the process is challenging due to the need to prepare the paste in an aqueous-based solution with low binder content in order to keep the process environmentally friendly. For this purpose an existing single-extruder FEF machine was modified to a multiple-extruder machine capable of extruding multiple materials.

For this purpose an existing single-extruder FEF machine was modified to a multiple-extruder machine capable of extruding multiple materials. By using one of these extruders to deposit the sacrificial material, which is removed during the post-processing, this FEF process with multiple extruders is capable of building parts with different types of features, e.g., internal holes and overhangs. After identifying methyl cellulose as a workable sacrificial material, mathematical models representing the dynamics of paste extrusion were developed through extrusion experiments with alumina and methyl cellulose pastes. For demonstration the fabricated green part was a cube-shaped part with a square hole in each side. The green part was shown to have good dimensional accuracy with respect to the CAD model. The green part was freeze-dried, debinded and sintered to remove the sacrificial material and achieve a high-density ceramic part.2.

The green part was freeze-dried, debinded and sintered to remove the sacrificial material and achieve a high-density ceramic part.2. Experimental SetupThe experimental setup for the freeze-form extrusion fabrication (FEF) process with multiple extruders consists of a motion subsystem, a real-time control subsystem, and extrusion devices. Figure 1 shows a phto of the overall system. The system consists of three linear axes (Parker Hannifin) each driven by a stepper motor (Empire Magnetics). The X, Y and Z axes each have 254 mm of travel range. The gantry motors each have a stepping angle of 1.8o. Each motor has a resolver that measures the angular position and feeds the signal to a resolver-to-digital encoder converter module (RDE). The RDE converts the resolver signal to an equivalent encoder feedback signal, allowing the resolution of each axis at 2.5 μm per step.

The RDE converts the resolver signal to an equivalent encoder feedback signal, allowing the resolution of each axis at 2.5 μm per step. For each axis, themaximum angular velocity is 50 rev/s, which provides a maximum velocity of 250 mm/s in linear movement. The maximum motor acceleration is 50 rev/s2, which provides a maximum linear acceleration of 250 mm/s2. The drives (National Instruments, NI) are used to amplify power from the two motion control cards for the stepper motors. Each card has outputs for up to four stepper motors plus inputs from encoders and limit switches for the four motors. The four outputs control the three stepper motors for the gantry axes and one stepper motor for an extruder. Another NI drive is used to regulate the other two extruder motors and to receive inputs from encoders and limit switches for two axes. The signals are amplified and then sent through a 32-pin connector to the two motion control cards.

The signals are amplified and then sent through a 32-pin connector to the two motion control cards. A National Instruments PXI 8176 real-time system with LabVIEW is used for software development and graphical user interface. The NI PXI-6025 multifunction data acquisition card is used for data input and output.Figure 1. Experimental setup of the multiple-extruder FEF machineThree extruders could be used for the deposition of three different pastes. They have the resolution of 2,846 steps/mm and could reach 72,882 steps/mm if using micro-stepping. The lead screw of each extruder had 20 cm of linear travel. Three 50 cm3 syringes were used to contain the paste material for extrusion. A removable 580 μm diameter nozzle was used for depositing material. By adjusting the commanded speed of stepper motor, the material deposition rate and the force on the syringe for extrusion could be controlled.

By adjusting the commanded speed of stepper motor, the material deposition rate and the force on the syringe for extrusion could be controlled. Three LC-301 load cells from Omega Engineering were mounted on the three extruder rams for extrusion force feedback. The load cell measured the extrusion force in the range of 0 to 2,250 N while the ram was applying pressure to the ceramic paste.An air conditioning system was installed on the freezer box to maintain the work environment at sub-zero temperatures in the experimental setup. The syringes were mounted with heating cylinders to prevent the paste from freezing before exiting the nozzle during part fabrication. Two temperature controllers (Omega Engineering) were used to control the temperatures inside the box and in the heating cylinders.

Two temperature controllers (Omega Engineering) were used to control the temperatures inside the box and in the heating cylinders. The temperature inside the box was normally set to -10 °C, and the temperature for the heating cylinders was set to 20 °C to result in 5 °C temperature at the nozzles.Commercially available software, along with Matlab, was used with LabVIEW to coordinate the motion of the extruders and to switch between different materials during the fabrication process. Skeinforge, which is an open-source software package, was used to generate the tool path for both depositing both alumina and methyl cellulose pasts. The STL file of a CAD model is required to generate the tool path. The inputs to Skeinforge also include parameters such as extrudate width, stand-off distance, and table speed. The output of Skeinforge is tool path in G-code. A converter script in C++ was written to translate from G-code to an executable code by the FEF system.3.

The output of Skeinforge is tool path in G-code. A converter script in C++ was written to translate from G-code to an executable code by the FEF system.3. Identification of Sacrificial MaterialTwo different materials were tested to determine the suitable sacrificial material for the FEF process. The first sacrificial material tested was carbon black, based on the previous research by Lewis et al. [18]. The composition of the carbon black paste used in our experiment was 48 vol.% carbon black ink, 4 vol.% methyl cellulose binder and 48 vol.% water. The carbon black paste did not work well due to viscosity inconsistency. The carbon black reacted differently at room temperature, which had ~30% humidity, than inside the freezer box at a sub- zero temperature, which had ~80% humidity. The high humidity inside the freezer box made the carbon black paste smudged after deposition.

The high humidity inside the freezer box made the carbon black paste smudged after deposition. This phenomenon could be explained by the fact that the water content in the carbon black paste is increased by 4% when there is an increase of 10% in humidity in the environment [19]. Thus the carbon black paste viscosity was increased by almost 20 vol.% from its original composition due to increase of humidity from 30% to 80% when the paste was inside the freezer box. Carbon black was successfully used as a sacrificial material in the Direct Ink Writing (DIW) [18], which was performed at room temperature (without little humidity change) with the part fabricated in a pool of oil to prevent change of viscosity. However, it is not suitable for the FEF process because the FEF process is performed at a sub-zero temperature (with large humidity change).Therefore, an alternative sacrificial material was investigated due to the incompatibility of carbon black for the FEF process.

Methyl cellulose was identified as a favorable sacrificial material due to its good rheological properties at sub-zero temperatures. The methyl cellulose paste used in our study had 10 vol.% methyl cellulose and 90 vol.% water.4. Modeling of Paste Extrusion DynamicsThe extrusion of two different aqueous pastes, alumina and methyl cellulose, was modeled by running extrusion tests to determine the values of parameters of a first-order system in the form of a transfer function. These values were determined from the system’s response to areference input of ram velocity, with the extrusion force as the output. The transfer function between the extrusion force and the ram velocity is given by()()1FsKVss(1)where s is the Laplace operator, F is the extrusion force, V is the ram velocity, K is the extrusion process gain, and τ is the extrusion process time constant.

The time constant for the first-order model for paste extrusion is calculated using the following equation2.2rT(2)where Tr is the rise time (time taken from 10% to 90% of steady-state force). The gain is calculated usingssFKV(3)where Fss is the steady-state force.Table 1 shows the parameters calculated from the data collected from the experiment. It can be seen from this table that the time constant decreases as the ram velocity increases. By comparing the time constant of 22.15 s for the ram velocity of 0.1 mm/s and the time constant of 1.39 s for the ram velocity of 1 mm/s, we see that there is one order of magnitude difference in the time constant. The data from Table 1 was used to calculate the response of a first-order system using Equation (1). Figure 2 compares the simulation (using the calculated time constants) and experimental results, indicating that a first-order model is a good approximation of aluminum paste extrusion dynamics.

Figure 2 compares the simulation (using the calculated time constants) and experimental results, indicating that a first-order model is a good approximation of aluminum paste extrusion dynamics. From the values in Table 1, the relationship between the steady-state extrusion force vs. ram velocity is given in Figure 3, which shows a non-linear relationship between the steady-state force output and the ram velocity input. This relationship is attributed to a nonlinear relationship between the paste viscosity and the shear rate in a non- Newtonian fluid described by the Herschel –Bulkley model [20].Table 1. Time constant, steady-state force, and gain for alumina paste extrusion at different ram velocity inputsVelocity (mm/s)T (s)Fss (N)K (N/mm/s)0.122.15216.882168.80.27.20280.101400.50.34.43301.471004.90.53.20340.27680.540.71.76363.21518.8711.39405.25405.25Figure 2.

Comparison of simulation and experimental results based on gain and time constant for alumina paste extrusion from Table 1Figure 3. Relationship between steady-state force and ram velocity for alumina pasteFigure 4 shows the relationship between the time constant and the ram velocity for alumina paste extrusion using the parameter values in Table 1. The nonlinear relationship between time constant and ram velocity can be explained using the dynamic modeling of paste behavior by Li et al. [21,22], which considered some volume of air trapped inside the syringe and derived a dynamic extrusion force model to show that the time constant decreases when the extrusion force or ram velocity increases.Figure 4.

Relationship between time constant and ram velocity for alumina paste extrusionIn order to verify the values obtained in Table 1 and to apply a feedback force controller, the first-order model parameters were determined from another experiment where the ram velocity input was stepped alternately between 1 and -1 mm/s as shown in Figure 5. The time constant and system gain parameters were calculated in the digital domain using the Recursive Least Square method (RLS). A first-order dynamic model describing the paste extrusion force of the process in the digital domain is()(1)()()FzbKaGzVzzaza(4)where z is the forward shift operator, F is the extrusion force (N), V is the command velocity (mm/s), and K is the unknown model gain (N/mm/s).The time constant and gain, respectively, are()Tlna(5)1bKa(6)where T is the sampling time, and τ is the time constant.Figure 5.

Comparison of modeled and measured results with a step reference input alternating between 1 and -1 mm/s for extrusion of alumina pasteThe extrusion of methyl cellulose paste also exhibits a first-order behavior between the extrusion force output and the input ram velocity. Based on the obtained force-velocity data, the time constant and steady-state force were calculated and given in Table 2. Like the extrusion of alumina paste, the time constant decreases with increase in ram velocity in the extrusion of methyl cellulose paste.Table 2. Time constant, steady-state force, and gain for methyl cellulose paste at different ram velocity inputsVelocity (mm/s)T (s)Fss (N)K (N/mm/s)0.118.55176.231762.30.26.10222.141110.70.33.44248.37827.90.52.90296.24592.480.71.75320.54457.1411.36353.20353.205.

Controller Design and ImplementationDue to certain effects of extrusion such as the release of air bubbles, breakdown of agglomerates, and change in paste properties as a result of liquid phase migration, the extrusion force is difficult to control in an open loop [16-17]. Thus a General Tracking Controller (GTC) was implemented in this study to regulate the extrusion force in a close-loop manner and to allow extrusion on demand in the FEF process. The objectives of the closed-loop controller are to extrude the paste at a constant rate, to coordinate the start of extrusion with the gantry motion, and to reject disturbances with desired error dynamics. The block diagram of a GTC controller is depicted in Figure 6.Figure 6. Block diagram of the general tracking controllerThe control signal is related to error and reference signals by()()()()()()()()bzvzVzvzazRzgzEz(7)()()vzazwhere disturbance for a step input.

We achieved this by making the second pole at least one order of magnitude smaller than the dominant pole, and shortening the settling time as much as possible without causing system instability with the dominant pole.A general tracking controller as described above was implemented for control of alumina paste extrusion. Using Equations (11) – (14) and the model parameters calculated with the RLS method, the obtained parameters were = 1.06. To validate the working of the = -1.07 and GTC, an input square signal was given with low and high limits of 100 N and 400 N, respectively, as shown in Figure 7. Two over-damped poles were selected with the time constants τ1= 0.5 s and τ2= 0.05 s, thus the time constant of the closed-loop system was determined by τ1. The time constant τ1 was determined by the mechanical limitations of the FEF system.

The time constant τ1 was determined by the mechanical limitations of the FEF system. The maximum input velocity for the alumina paste extrusion process in the FEF system was 2 mm/s, at which the maximum force of 550 N provided by the stepper motor in the system was reached. The open-loop time constant for 2 mm/s ram velocity input was calculated using Equation (4) and the time constant obtained was τ = 0.48 s . The closed-loop time constant was chosen as 0.5 s, in order to have the quickest response of the system without overloading the motor. The reference force for extrusion of alumina paste was set at 400 N, which is slightly less than 550 N for margin of safety. The force of 100 N was used to stop extrusion based on experimental test runs. The GTC tracked the reference force with an error of +/-10 N, indicating a consistent extrusion rate. From Figure 8, the time constant of the closed-loop system calculated from the rise time is approximately 0.5 s using Equation (2).

From Figure 8, the time constant of the closed-loop system calculated from the rise time is approximately 0.5 s using Equation (2). This time constant is reduced by 65% from 1.39 s (Table 1) in the open-loop system for a 400 N output force (or 1 mm/s in ram velocity). For a smaller extrusion force (or ram velocity), the percentage reduction in the time constant is larger because the time constant in the closed-loop system remains unchanged while the time constant in the open-loop system increases with decrease in extrusion force (or ram velocity) as seen in Table 1. The ram input velocity was set to +/- 2mm/s in order to have a faster response of the system without getting into instability. Higher velocities would cause the stepper motor to skip and not able to track the reference force as desired.1g0gFigure 7. Response of the general tracking controller for a reference ram force input for extrusion of alumina pasteFigure 8. Reference vs.

Response of the general tracking controller for a reference ram force input for extrusion of alumina pasteFigure 8. Reference vs. measured force using the GTC with a rise time of 1s for extrusion of alumina pasteA series of tests was conducted to find a relationship between the extrudate velocity and the reference extrusion force in the close-loop system. Table 3 was generated by varying the reference force from 150 N to 200 N, then to 300 N, and finally to 400 N. The nozzle tip was positioned along the Z direction 50 mm above the X-Y table in the gantry system. The extrudate velocity was calculated by measuring the time it took for the paste to reach the X-Y table from the moment it was extruded from the tip. A stopwatch with a resolution of 0.1 s was used for this time measurement. Six test runs per reference force were conducted and the measured data were averaged.

A stopwatch with a resolution of 0.1 s was used for this time measurement. Six test runs per reference force were conducted and the measured data were averaged. The different tests per reference force show similar values, verifying the advantage of using the GTC for consistent paste extrusion.Table 3. Relationship between reference extrusion force and extrudate velocity for alumina pasteTest runDistance(mm)Force(N)Test 1 (s)Test 2 (s)Test 3 (s)Test 4 (s)Test 5 (s)Test 6 (s)Mean extrudate speed (mm/s)15015080.182.385.484.382.182.80.625020042.140.144.243.242.142.31.1835030020.121.222.021.820.921.22.3545040012.313.114.213.512.913.23.78The relationship between extrusion force and extrudate velocity in Table 3 is plotted in Figure 9.

The y intercept in this figure shows that when the extrudate velocity is zero, the steady- state force is 106.6 N, indicating that the minimum force required to extrude the alumina paste from the nozzle tip is slightly larger than 100 N. If the extrusion force is less than this threshold, extrusion of the paste will cease because the applied ram force is not sufficient to overcome the shear stress required for the paste flow.Figure 9. Reference force extrusion vs. extrudate velocity for alumina pasteFor extrusion of methyl cellulose paste, a general tracking controller was also implemented. A series of tests was conducted to find the relationship between the extrudate velocity and the reference force in the closed-loop system for extrusion of methyl cellulose paste. Based on experimental test runs, the reference force of 350 N was used to start the paste extrusion and the reference value of 50 N was used to stop the extrusion.

Based on experimental test runs, the reference force of 350 N was used to start the paste extrusion and the reference value of 50 N was used to stop the extrusion. The closed-loop time constant was calculated based on the open-loop time constant for a ram velocity of 2 mm/s. The time constant of 0.5 s was again chosen, which is the same as the time constant used in theclosed-loop control of alumina paste extrusion. With this time constant the rise time was approximately 1 s, and thus the time constant of the open-loop system (1.36 s) was reduced to 0.5 s (65% reduction) for the steady-state extrusion force of 350 N in the closed-loop system. The time constant of the closed-loop system was approximately 0.5 s as expected. The small time constant and force error result in a consistent extrusion rate when depositing methyl cellulose paste during part fabrication.6.

The small time constant and force error result in a consistent extrusion rate when depositing methyl cellulose paste during part fabrication.6. Effects of Process ParametersThe quality of part fabricated by the FEF process is affected by parameters including the stand-off distance and table speed. This section discusses the determination of effective parameter values for the FEF process with the use of sacrificial material. In all of the experiments, the diameter of the needle nozzle was 580 µm.By using the maximum extrusion force in the FEF process, which was 400 N for alumina and 350 N for methyl cellulose, the extrudate diameter was determined experimentally as 780 µm and 760 µm, respectively, for a freezer temperature of -10 °C. Experiments were conducted to measure the extrudate width by varying the stand-off distance. The range of stand-off distance varied from 700 µm to 400 µm in decrements of 100 µm in the experiments.

Experiments were conducted to measure the extrudate width by varying the stand-off distance. The range of stand-off distance varied from 700 µm to 400 µm in decrements of 100 µm in the experiments. The path was a serpentine trajectory of 4 lines and the table speed was 8 mm/s. The lines were measured using a caliper with a resolution of 10 µm. Five lines were deposited for each stand-off distance in order to see the repeatability of the process in having a consistent extrudate width. It was observed that the shorter the stand-off distance, the wider the extrudate line. Note that the extrudate diameter of the paste coming out of the nozzle had a diameter larger than the stand-off distance. This caused the extrudate line to flatten out, resulting in an approximately rectangular (instead of circular) cross-sectional area. The extrudate widths at different stand-off distances for alumina paste extrusion are given in Table 4.

The extrudate widths at different stand-off distances for alumina paste extrusion are given in Table 4. The same experiment was repeated for methyl cellulose paste.Table 4. Relationship between stand-off distance and extrudate width for alumina pasteTest runStand-off distance(µm)Line 1Width (mm)Line 2Width (mm)Line 3Width (mm)Line 4Width (mm)Line 5Width (mm)Mean Width (mm)14001.231.251.221.211.231.2225001.151.121.101.111.111.1136001.00.980.960.950.940.9647000.800.790.810.800.760.79The cross sections from rectangular blocks fabricated at different stand-off distances are shown in Figure 10. The block with 600 µm stand-off distance (Figure 10a) shows extrudate lines with rounder edges and some holes between the extrudate lines.

The block with 600 µm stand-off distance (Figure 10a) shows extrudate lines with rounder edges and some holes between the extrudate lines. The blocks built at 500 µm and 400 µm stand-off distances (Figure 10b & 10c) show extrudate lines with sharper edges and also fewer and smaller holes between the extrudate lines. Therefore, 400 µm and 500 µm are more desirable stand-off distances. In further experiments the stand-off distance of 500 µm was used in extruding both alumina and methyl cellulose pastes.Figure 10. Cross sections of extrudate blocks with the stand-off distance of (a) 600 µm, (b) 500 µm, and (c) 400 µmThe table speed is crucial to the quality paste deposition. Six test runs were conducted in which the table speed was varied from 4 mm/s to 14 mm/s in increments of 2 mm/s, while the extrudate speed was 4 mm/s.

Six test runs were conducted in which the table speed was varied from 4 mm/s to 14 mm/s in increments of 2 mm/s, while the extrudate speed was 4 mm/s. Figure 11 and shows the results of extrusion using an extrusion force of 400 N for alumina paste and 350 N for methyl cellulose paste (at 500 µm stand-off distance). It can be seen that when the table speed was 8 mm/s, the deposition of both alumina and methyl cellulose pastes had relatively uniform width compared with the other table speeds. At velocities higher than 10 mm/s, inconsistency in the width of extrudate line was observed. Such a deposition could lead to gaps between lines during the rastering process, resulting in poor-quality parts. It was reported previously that maintaining the extrusion speed at the table speed is advisable [16].

It was reported previously that maintaining the extrusion speed at the table speed is advisable [16]. However, having the table speed same as the extrudate speed could possibly result in an over-extrusion of paste surrounding the nozzle tip due to the stand-off distance less than the extrudate diameter, if the nozzle wall is not thick enough.Figure 11. Different table speeds for extrusion of alumina paste (left) at 400 N extrusion force and for extrusion of methyl cellulose paste (right) at 350 N extrusion forceThe equation based on conservation of paste flow is24exTVDVwh(17)where VT is the Table speed, Vex is the extrudate speed, D is the extrudate diameter (780 µm for alumina), w is the extrudate width, and h is the stand-off distance. From Equation (17) it can be calculated that when the table speed is 4 mm/s, the extrudate width is 1.91 mm.

From Equation (17) it can be calculated that when the table speed is 4 mm/s, the extrudate width is 1.91 mm. This is much larger than the outer diameter of the nozzle, thus at this speed there was an over-extrusion ofpaste material and the width of the extruded paste line was not very uniform in Fig. 11. When the table speed is 6 mm/s, the line width of the extruded paste is 1.21 mm, which is slightly larger than the nozzle’s outer diameter. When the table speed is 8 mm/s, the extruded paste line width is 0.95 mm, which is slightly smaller than the nozzle’s outer diameter. The above explains why the extrudate line is most uniform in Fig. 11 when the table speed is 8 mm/s.7. Part Fabrication Results and DiscussionSeveral parts were fabricated to demonstrate the capabilities of the developed FEF system. The freezer temperature was -10 °C, and the stand-off distance used was 500 µm. For purpose of improving part quality, there was a waiting time of 10 s between layers.

The freezer temperature was -10 °C, and the stand-off distance used was 500 µm. For purpose of improving part quality, there was a waiting time of 10 s between layers. Figure 12 shows the CAD model for a cube with square through holes in all sides and an alumina part fabricated by the FEF process with use of sacrificial material for the CAD model.Figure 12. CAD model of a cube with a square hole on each side and the extruded part with sacrificial materialThe sacrificial material methyl cellulose was burnt out to remove it from the part prior to sintering. Then sintering was performed to densify the alumina part into a solid ceramic structure. The binder burnout process began slowly with a ramp of 2 °C/min and continued until reaching 400 °C. After the methyl cellulose was burned out, the slope of ramp was increased to a rate of 10° C/min until 1550 °C. After reaching 1550 °C, the temperature was kept constant for one hour to sinter the part.

After reaching 1550 °C, the temperature was kept constant for one hour to sinter the part. Then the furnace was cooled down at a ramp of 75 °C/min until reaching room temperature. The slow ramp rate of 2 °C/min to 400 °C was based on previous experiments in burning out organics from ceramic bodies [16]. The critical temperatures in the binder removal and part sintering were obtained using Thermogravimetric Analysis (TGA) to measure the weight loss of material as a function of temperature.Dimensional measurements were taken from the green parts using a caliper with a resolution of 1 µm. The cube and mushroom parts were selected because their dimensions could be measured easily with a caliper. In Figure 12, each measured dimension is labelled with a letter. Four measurements were taken from random locations from the green part. The measured data together with the means and standard deviations are given in Table 5.

Four measurements were taken from random locations from the green part. The measured data together with the means and standard deviations are given in Table 5. From this table, the dimensions of the green part for the cube varied between 0.1% to 3.0% compared with the dimensions of the CAD model. The differences were small and would suggest good accuracy of part fabricated by the FEF process.Table 5.

The differences were small and would suggest good accuracy of part fabricated by the FEF process.Table 5. Measurements for a cube-shaped part in its green stateMeasurement L(mm) W(mm) H(mm) A(mm) B(mm) C(mm) D(mm)CAD model30303010101010130.2030.1030.1710.1010.2010.611.0229.8129.8829.879.909.600.979.80329.7530.4030.239.809.800.909.20430.4029.9529.8011.001.100.969.90Mean30.0430.0830.0110.2010.109.709.90Difference0.040.080.10.200.100.300.10% Error0.100.200.302.01.03.01.0Standard Deviation0.310.230.210.050.610.600.70After sintering, eight measurements were taken from different points for each dimension for the cube-shaped part. Table 6 gives the measured data together with means and standard deviations for the sintered part.

Table 6 gives the measured data together with means and standard deviations for the sintered part. This table shows that the standard deviation varied between 0.7% and 1%, and that the percentage of shrinkage for the various dimensions was in the range of 16- 23% compared to the green part after sintering.Table 6.

Measurements for the cube-shaped part after sinteringTestL (mm)W (mm)H (mm)A (mm)B (mm)C (mm)D (mm)125.1925.1125.097.967.747.877.77225.0925.0724.887.907.657.807.70324.9225.0225.127.847.827.747.82425.3024.8724.978.047.807.907.86525.1524.9225.007.847.827.927.84625.0724.8724.777.947.767.887.92724.9625.1325.067.907.907.797.84825.8225.0824.808.007.827.917.78Mean25.18 25.00 24.96 7.927.787.857.81Standard deviation0.280.100.130.071 0.073 0.060.06% Shrinkage16171722231921It should be noted that part accuracy can be improved by using nozzles of smaller diameters. For example, a nozzle of 300 µm in diameter can be used in lieu of the nozzle of 580 µm used in the current FEF process. However, using a smaller nozzle will increase the part build time.

For example, a nozzle of 300 µm in diameter can be used in lieu of the nozzle of 580 µm used in the current FEF process. However, using a smaller nozzle will increase the part build time. To decrease the building time, the extrusion force can be increased in order to increase the extrudate velocity. However, since the extrusion force is limited by the motor power, more powerful motors will be needed in order to increase the extrusion force.8. ConclusionsA freeze-form extrusion fabrication (FEF) system with multiple extruders operating in a freezing environment (about -10 °C) has been developed and successfully implemented. The parts were built by extruding and depositing aqueous alumina paste as the main material, and methyl cellulose was identified as the sacrificial material. An empirical first-order model was used to represent the extrusion dynamics for both alumina and methyl cellulose pastes, with the ram velocity as the input and extrusion force as the output.

An empirical first-order model was used to represent the extrusion dynamics for both alumina and methyl cellulose pastes, with the ram velocity as the input and extrusion force as the output. A general tracking controller was applied to control the extrusion force. The controller reduced the time constant for the closed- loop system by 65% and more when compared to the open-loop control system in extrusion of both alumina and methyl cellulose pastes. The capability of the developed system was demonstrated by fabricating a cube-shaped part with square holes on all sides. The dimensional accuracy of the green part was in the range of 0.1 – 3.0% compared to the CAD model. The sacrificial material was removed successfully during the binder burnout.

The dimensional accuracy of the green part was in the range of 0.1 – 3.0% compared to the CAD model. The sacrificial material was removed successfully during the binder burnout. Shrinkage varied between 16% and 23% for the various dimensions of the part after sintering.AcknowledgementsThis project was funded by NSF grant #CMMI-0856419 with matching support from the Boeing company through the Center for Aerospace Manufacturing Technologies at the Missouri University of Science and Technology.References[1]“Advanced Ceramics Technology Roadmap-Charting Our Course,” Sponsored by U.S. Advanced Ceramic Association and U.S Department of Energy, Prepared by Energetics, Inc. and Richerson and Associates, December 2000.[2] S. Rangarajan, Q. Qi, N. Venkataraman, A. Safari, and S. C. Danforth, “Powder Processing, Rheology, and Mechanical Properties of Feedstock for Fused Deposition of Si3N4 Ceramics,” Journal of the American Ceramic Society, 83(7), pp. 1663-1669, 2000.[3] A.

Extrusion deposition is especially advantageous since multiple nozzles easily accommodate the inclusion of additional materials. However, discrete interfaces and sudden composition changes can limit the functionality of a printed part through inherently weak bonding. Furthermore, same-layer transitions are not only difficult to execute, but also further amplify structural weaknesses by creating multiple discrete interfaces. Therefore, successfully implementing a blended, continuous gradient will greatly advance the applicability of FGM in additive manufacturing. The pellet-fed nature and integrated screw design of the Big Area Additive Manufacturing system enables material mixing needed for development of this capability.

The pellet-fed nature and integrated screw design of the Big Area Additive Manufacturing system enables material mixing needed for development of this capability. Using constituent content analysis, this study evaluates the transition behavior of a neat ABS/CF-ABS material pair and characterizes the repeatability of the mixing and printing process, which ultimately leads to control of site-specific material deposition and properties.IntroductionAdditive manufacturing (AM) capabilities have been continuously expanded in recentyears, and its potential for industrial applications has maintained a high level of interest. One of several areas being fervently explored is the adapting current processes to produce Multi- Material (MM) and Functionally Graded Material (FGM) structures. While many AM systems were developed for single-material use, Multi-Material Additive Manufacturing (MMAM) provides increased functionality.

While many AM systems were developed for single-material use, Multi-Material Additive Manufacturing (MMAM) provides increased functionality. A process is generally considered to truly be MMAM only if it utilizes more than one material and does not require pre-mixing, pre-compositing, or non-AM post-processing treatment [1]. On the contrary, an FGM can be constructed through one or more materials and is most accurately characterized by a spatial variation in composition or densification to achieve a desired functional performance [2]. Within this definition lie two classifications of FGM’s: stepwise and continuous. A stepwise FGM relies on a multi-layered approach with discrete interfaces appearing periodically through the structure to create an overarching structure change via regions of differing microstructures.

A continuous FGM, however, exhibits a constant change in microstructural composition across the occupied spatial region [3].1793MMAM has been explored in recent years with a variety of processes and methods, butthe majority still tend to produce MM structures by assigning different materials to different layers. This approach experiences issues associated with the difficulty of bonding dissimilar materials. For example, production of a sandwich structure with a Stratasys PolyJet Connex 350 system utilizing an adhesives-based approach and T-Peel geometry found that fracture consistently occurred at material interfaces [4]. Although further investigation demonstrated that aligning the material interface perpendicular to build direction improved fracture resistance along the interface, it did not completely eliminate the interface vulnerability [5].

Similarly, Brischetto et al investigated sandwich structure configurations produced with a multi-nozzle Fused Filament Fabrication (FFF) Sharebot NG 3D printer. Their work concluded that a fully single-material PolyLactic Acid (PLA) structure outperformed one that substituted ABS for the external face sheets when suing a honeycomb core. Implementing a solid PLA core improved performance, leading them to suggest both poor adhesion and premature cooling due to nozzle switching is to blame [6]. Roger et al explored three techniques for improving the interface strength between differently-colored ABS, and while an interpenetrating pattern where each layer’s interface boundary is shifted from one side to another did improve ultimate tensile strength, the configuration still significantly underperformed compared to single-material constructions [7]. Another approach assessed the influence of structural arrangement of the ABS and PLA regions on mechanical performance.

Another approach assessed the influence of structural arrangement of the ABS and PLA regions on mechanical performance. Kim et al found that incorporating material changes in both the print direction and the z-direction dramatically improved performance and assisted to limit the detrimental influence of void formation between dissimilar beads [8].Application of MMAM technology to incorporate FGM into AM has also been prevalent.Vu et al further expanded their MMAM T-peel investigations to incorporate a stepwise FGM, which exhibited a significant increase in peel force and apparent fracture energy compared to a simple sandwich structure geometry [5]. Stratasys’s PolyJet technology was also implemented in construction of a soft, flexible robot shell to create an FGM that varied in elastic modulus from 1 MPa to 1 GPa, which allowed it to soften impacts while maintaining a semi-rigid overall shape and survive longer than its purely rigid or flexible counterparts [9].

Stratasys has also developed partnerships intended to attack the problem from the software side. Ezair et al developed a software interface that allows designers to volumetrically apply material choices to their designs instead of relying on contemporary boundary-representation techniques. With this framework, the team produced FGM structures that demonstrated control of porous regions and deposition of differing materials via repeatable patterns [10]. Investigation of FGM has even been conducted with large-scale AM. In one such study, Cincinnati’s Big Area Additive Manufacturing (BAAM) system was used to print two different types of stepwise FGM sandwich structures, which demonstrated the potential for FGM implementation to impart both cost and weight savings even within AM [11].The goal of this investigation is to expand upon the initial BAAM work and investigate the implantation of a continuous FGM in large-scale AM.

In particular, it will characterize the continuous transition produced by in-situ material switching during printing. The BAAM at Oak Ridge National Laboratory’s Manufacturing Demonstration Facility provides a unique opportunity for this study due to recent alterations enabling a dual-hopper configuration while printing, which allows for on-the-fly material switching made possible by the pelletized feedstock and extrusion system. This set-up provides significant advantages over multi-nozzle1794systems given the inherent blending of a screw-based extruder and the lack of stoppage during material switching. For this investigation, carbon fiber reinforced ABS and neat ABS are chosen as the bi-material system. Maintaining the same polymer matrix between the two materials allows for the printing process to be characterized with minimal outside factors.ExperimentalTwo grades of ABS were obtained in a pelletized form from Techmer Engineered Solutions.

The first was a Neat ABS (HIFILL ABS 1000 3DP)) and the second was a 20 wt. % carbon fiber ABS (Electrafil J-1200/CF/20 3DP). Both materials were used in printing on the Manufacturing Demonstration Facility’s (MDF) Big Area Additive Manufacturing (BAAM) – B that features an innovative dual-hopper system for multi-material printing. Seen in Figure 1, the system functions on a rocker mechanism that forces one hopper’s feeding tube closed and off the barrel while simultaneously opening and moving the other into position. The materials were dried at 85°C overnight prior to printing.Figure 1: The newly installed dual-hopper system on the MDF's BAAM-B. This system enables in-situ material switching for MM and FGM large-scale printing of thermoplastics.Printing Plan and ParametersTo demonstrate a complete material transition from in-situ feedstock switching, thesingle-walled rectangle geometry in Figure 2A was chosen.

By implementing a hopper switch twice within a layer at opposing corners, a complete transition was confined to the long sides of a rectangle with dimensions 11.5 ft by 5 ft. The printing process for a single print consisted of three stages: an initial single-layer that was printed and removed, a 20-layer rectangle that was printed and removed, then a final single-layer was printed and removed. No purging was1795A.)B.)4Side 4-3 (Neat ABS)3(/)a: - l ,-., ~ -11).j:::,,Vl0nSwitch Point/,-., VJ~ 0 ....VJ~ --µ.. u' - ' ("')Start/Switch PointFigure 2: A.) Schematic of the printed rectangle. The expected composition of each side is indicating with labels, and the corners marked 1 and 3 label the two points within a layer that a material switch was initiated. B.) The 4 locations along each side where samples were extracted.conducted between stages.

B.) The 4 locations along each side where samples were extracted.conducted between stages. The purpose of the 20-layer stage is merely to incorporate numerous in-situ switches over the lifecycle of a print so that the effects of continual switching on the transition in a later layer can be observed. In total, three prints were conducted and analyzed to identify the consistency of the material switch, the repeatability of the transition throughout the layers of a print, and the repeatability of the transition in after multiple prints. Furthermore, the impact of transition direction (Material A (cid:198) Material B vs. Material B (cid:198) Material A) will be evident by comparing opposing sides of the rectangle.These prints utilized a nozzle of 10.16 mm (0.4 in) diameter with a non-mixing screw.

Material B (cid:198) Material A) will be evident by comparing opposing sides of the rectangle.These prints utilized a nozzle of 10.16 mm (0.4 in) diameter with a non-mixing screw. The print parameters maintained a constant screw speed of 300 RPM with travel speed of 106.7 mm (4.2 in/s) with layer height 6.35 mm (0.25 in), bead width 15.24 mm (0.6 in), melt temperature of 245°C, and bed temperature of 100°C.Constituent Content AnalysisTracking the transition from one material to the next requires determination of materialcomposition at any given point. For this material pair, wt % carbon fiber is used to track the change between materials. ASTM D3171 Procedure H: Carbonization in Nitrogen (CIN) protects the carbon fibers with an inert nitrogen atmosphere while eliminating the polymer matrix [12]. Additionally, it implements a 100% Neat ABS control to ensure any remaining debris from polymer burn-off is considered.

Additionally, it implements a 100% Neat ABS control to ensure any remaining debris from polymer burn-off is considered. A Lindberg Blue M nitrogen tube furnace was used to conduct Procedure H. The remaining percent carbon fiber is compared to the nominal constituent content of the 20% CF/ABS, which was found to be closer to 21 wt % carbon fiber. All mass measurements were taken with Mettler Toledo Balance, XS204 Deltarange to a 0.1 mg accuracy.1796To ensure accurate data, each single-layer rectangle was sectioned into 152.4 mm (6 in) rods and labeled to indicate the order and direction of the print. Using an Isomet 1000 Precision Saw, samples 8.4 mm in length were taken from the ends of these rods at the desired locations, which were chosen so that samples from different layers and prints would maintain the same distance from most recent switch for direct comparison.

As can be seen in Figure 2B, each sample set consisted of four locations on each side chosen such that both neat and transition regions would be represented.CIN testing was divided into four categories to determine the consistency of thetransition, repeatability of the printing, repeatability of layers, and the construction of the transition curve. In the first set of experiments, three consecutive samples were taken to confirm that CIN yields the same fiber content for each one, thus testing for any variance in the procedure or during a given transition. Variance between different prints of the same structure were then considered to ensure that a transition is predictable and repeatable. In the third experimental set, the first and final layers are compared to establish if continuous switching throughout a print causes a change in the transition behavior.

In the third experimental set, the first and final layers are compared to establish if continuous switching throughout a print causes a change in the transition behavior. Finally, numerous additional samples were taken from a single print to construct a beginning to end transition curve that describes the switch from CF/ABS to Neat ABS and back.Results and DiscussionRepeatability of Sampling LocationFigure 3 shows the CIN results of subsequent samples at four locations for both transitiondirections. Three locations required additional testing, for while they showed consistency in measurement, there was an issue with the control sample that initially returned negative values. The repeat testing utilized only one sample to obtain a reasonable fiber measurement as the original measurements had already proven consistency.

The repeat testing utilized only one sample to obtain a reasonable fiber measurement as the original measurements had already proven consistency. Additionally, the 1.98 m location in Figure 3A initially showed a significant difference in the three samples from that spot, but repeat testing confirmed that the two of the measurements were anomalies. Thus, Figure 3A and 3B demonstrate that the material transition is consistent.

Thus, Figure 3A and 3B demonstrate that the material transition is consistent. Moving forward, it allows for the treatment of single 8.4 mm sample as representative of the surrounding area.A.)Consistency Through a Bead: CF/ABS ➔ABSB.)Consistency Through a Bead: ABS ➔ CF/ABSI "A" Samples ■ "B" Samples ■ "C- Samples25%25%°"1.681.983.35Distance from Hopper Switch (m)4.8820% I IS% 11 ~ 10% _ge s" °"1.682593.35Distance from Hopper Switch (m)Figure 3: The carbon fiber content of three consecutive 8.4 mm samples taken from four locations along the A.) CF/ABS to ABS transition and B.) the ABS to CF/ABS transition.17974.88Repeatability of the PrintingFigure 4 shows the resulting fiber content as measured by CIN for each of the three prints in both transition directions. For both directions and all locations, the variation in measured fiber content is minimal.

For both directions and all locations, the variation in measured fiber content is minimal. The maximum fiber wt % difference observed between measurements is less than 1.5 wt %, and the largest standard deviation at a location was 0.66%.

This indicates the in- situ switching process can be reliably charted and replicated between prints, which will allow for accurate modelling and application of a shape function to the printing process.A.)Location Comparison Across Prints: CF/A BS➔ABSB .)Location Comparison Across Print s: ABS➔CF/ABS■ Print 1 ■ Print 2 ■ Print 3■ Print 1 ■ Print 2 ■ Print 325%25%i: 20% 3 j 15%u: r:0%1.681.98 ---3.354.880%1.682.593.35Distance from Hopper Switch (m)Distance from Hopper Switch (m)Figure 4: The carbon fiber content at four specified distances from where the hopper switch was initiated for the A.) CF/ABS to ABS transition and B.) the ABS to CF/ABS transition.Repeatability of LayersFigure 5 contains the fiber content measured via CIN for layers 1 and 20 in both transition directions at four locations each.

Figure 5a shows consistent fiber content for locations outside of the transition, however, the location within the transitioning region consistently illustrated a difference between the layers. Fiber content is lower at later layers, implying that the transition region will drift or shift to occur faster or sooner in later layers compared to initial layers. In contrast, transitioning from a neat ABS to the carbon fiber filled ABS maintained a consistent fiber content at all spots. Thus, transition drift appears to be transition direction dependent and is heavily influenced by the presence of fiber fillers. A possible explanation is the build-up of residual fibers inside the barrel of the extruder assisting in a rapid change to CF/ABS. For example, even after fiber content drops below 1% and reaches steady state, the neat ABS portion still shows a visible fiber presence, as seen in Figure 5b.

u05%0%1.682.593.35Distance from Hopper Switch (m)Figure 5: The average fiber content of the three prints at each of the four locations for both the first and the last layer printed. A.) The CF/ABS to ABS transition and B.) The ABS to CF/ABS transition.Transition CurvesEstablishing confidence in the in-situ switching print capability as a repeatable process provided the needed support for analysis of a single print’s transitions as representative of the material pair’s behavior. Thus, additional data points taken from locations before, during, and after the transition were utilized to map the change in fiber content through the printed bead. Figure 6 shows the change in fiber content over distance for the transition from 20 wt % CF/ABS to Neat ABS. Due to the behavior it exhibited, the transition was divided into three distinct regions.

Figure 6 shows the change in fiber content over distance for the transition from 20 wt % CF/ABS to Neat ABS. Due to the behavior it exhibited, the transition was divided into three distinct regions. The first, the “Purge Zone”, consists of the first material due to the time it takes for the initial traces of the new material to travel from the top of the extruder through the barrel and to the nozzle. At this point, it enters the “Transition Zone,” where the carbon fiber content continuously changes until reaching a stable level associated with the second material, the “Steady State Zone.” In comparison to Figure 7, it is evident that the directionality of the transition imparts different characteristics into the transition.

Figure 8 combines the two graphs to more easily illustrate that the two transitions are not simply reversible, implying that the direction of the transition should be included as another parameter.17994.88ABS to CF/A BS Tra nsitions-ABS t o CF25%051015Distance from Hopper Switch (s) 25203035404550Purge ZoneTransitionSteady State20%Zone/l, 15% oi .c u:: C: 0 10% .c '" u 5%0%00.511.522.533.544.55Distance from Hopper Switch (m)Figure 6: The change in fiber content over time (s) and distance(m) from Neat ABS to 20 wt % CF/ABS.

More measurements were taken in the transition zone for improved resolution.CF/ABS to ABS Tra nsitions- CF to ABS25% O51015Distance from Hopper Switch (s) 25203035404550Purge ZoneTransition ZoneSteady State20%l, 15% QJ .c u::C: 0 10% .c '" u5%0%00.51.522.533.544.55Distance from Hopper Switch (m)Figure 7: The change in fiber content over time (s) and distance(m) from 20 wt % CF/ABS to Neat ABS. More measurements were taken in the transition zone for improved resolution.1800CF/ABS - ABS Transitions+- CF to ABS -+-ABS to CF25%051015Distance from Hopper Switch (s) 2520303540455020%'ii ~ 15% ~ ::! u..

C:_g 10% iii u5%0%00.511.522.533.544.55Distance fro m Hopper Swit ch (m)Figure 8: The CF/ABS to ABS and ABS to CF/ABS transitions overlaid on the same chart, highlighting the differences in transition behavior.Table 1 quantifies the zones shown in Figures 6 and 7. While CF/ABS to Neat ABSdemonstrates a shorter purge zone and quicker time to steady-state, the actual transition zone is slightly larger than that of the Neat ABS to CF/ABS direction. Overall, the CF/ABS to Neat ABS is a quicker process, but for both, the presence of a significant purge zone is an important distinction.

Overall, the CF/ABS to Neat ABS is a quicker process, but for both, the presence of a significant purge zone is an important distinction. For print design, it must be considered when inputting switch commands to not only ensure that steady state for Material B is reached when needed but also that the transition zone’s location is properly controlled.Table 1: Numerical comparison of Purge and Transition zonesUnitDirectionPurge ZoneTransition Zone CombinedMeterSecondCF/ABS- ABS ABS- CF/ABS CF/ABS-ABS ABS- CF/ABS1.83 2.29 17.14 21.431.37 1.22 12.86 11.433.20 3.51 30.00 32.86Conclusions and Future WorkImplementing an in-situ material change via feedstock switching for creation offunctionally graded and multi-material structures is viable. The transition was proven to be consistent through the bead with reliable, repeatable measurements of composition possible from just a 8.4 mm (1/3 in) samples.

The transition was proven to be consistent through the bead with reliable, repeatable measurements of composition possible from just a 8.4 mm (1/3 in) samples. Printing the same structure multiple times with in-situ switching1801produces near-identical transition behavior. However, evidence suggests that a material pair’s transition may not remain consistent after numerous switches within a single print since the CF/ABS to Neat ABS transition showed a significant difference between Layers 1 and 20 in the transition location. Conversely, transitioning in the opposite direction exhibited no change from the first layer to the twentieth, indicating that the change could be related to fiber build-up.

Conversely, transitioning in the opposite direction exhibited no change from the first layer to the twentieth, indicating that the change could be related to fiber build-up. Therefore, in-situ switching can, at the minimum, be considered a completely repeatable process for initial layers, so treatment of a single print as representative of a material pair’s transition behavior is justified.Creation of complete transition curves enabled a direct comparison of the opposite transition directions. Given that the transitions were not reversible, it indicates that the direction of a transition should also be considered a factor in predicting a material pair’s transition behavior. Furthermore, the complete curves illustrated the importance of considering the time it takes for a feedstock switch to manifest in the printing process.

Furthermore, the complete curves illustrated the importance of considering the time it takes for a feedstock switch to manifest in the printing process. Without proper characterization of the Purge Zone, initiating material switches within a print that accurately place the transition zone is not possible.This investigation is the beginning of a complete characterization of the dual-hopper’sFGM and MM capabilities. Moving forward, future studies will focus on quantifying the effects of various parameters. These include the directionality exhibited in this paper and what its cause is, the effect of differing screw designs and RPMS, and the influence of switching from one polymer matrix to another.AcknowledgementsResearch sponsored by the U.S. Department of Energy, Office of Energy Efficiency andRenewable Energy, Industrial Technologies Program, under contract DE-AC05-00OR22725 with UT-Battelle, LLC. Thank you to Cincinnati Incorporated for equipment and assistance.

Thank you to Cincinnati Incorporated for equipment and assistance. Special thanks to the University of Alabama Birmingham’s Materials Science Department for their help and cooperation with testing and data collection. This work was supported in part by Oak Ridge Institute for Science and Education through the Higher Education Research Experiences Program (HERE).References[1] M. Vaezi, S. Chianrabutra, B. Mellor, and S. Yang, "Multiple material additive[2][3]manufacturing – Part 1: a review," Virtual & Physical Prototyping, Article vol. 8, no. 1, pp. 19-50, 2013, doi: 10.1080/17452759.2013.778175. G. H. Loh, E. Pei, D. Harrison, and M. D. Monzón, "An overview of functionally graded additive manufacturing," Additive Manufacturing, vol. 23, pp. 34-44, 2018/10/01/ 2018, doi: https://doi.org/10.1016/j.addma.2018.06.023. G. Udupa, S. S. Rao, and K. V.

Vlastimil, "Expanding Material Property Space Maps with Functionally Graded Materials for Large Scale Additive Manufacturing," Solid Freeform Fabrication Symposium Proceedings, pp. 459-484, 2017. "ASTM D3171-15 Standard Test Methods for Constiuent Content of Composite Materials," ASTM International, West Conshohocken, PA., ASTM Standard 2015, doi: 10.1520/D3171-15.1803.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDEVELOPMENT OF HIGHLY FILLED BIO-BASED COMPOSITES FOR SUSTAINABLE, LOW-COST FEEDSTOCK: PROCESSING EFFECTS ON POROSITY AND FIBER ALIGNMENTKatie Copenhavera, Meghan Lamma, Amber HubbardaaManufacturing Science Division, Oak Ridge National Laboratory, Oak Ridge, TN, 37830.AbstractA poly(lactic acid) composite with a high loading of bio-based fibers was developed using a combination of high-aspect ratio (AR) wood pulp and low-AR wood flour along with viscosity modifiers to maximize mechanical performance, maintain processability, and lower the cost and embodied energy of the resulting feedstock.

An optimized composite formulation containing 40 wt.% of a blend of high- and low-AR natural fibers with a rice bran-based wax processing aid was scaled up to produce pellet feedstock using twin screw extrusion, and materials were compression and injection molded to investigate the effect of fiber alignment on material performance. The feedstock was then printed on the Big Area Additive Manufacturing system at Oak Ridge National Laboratory. Print parameters including temperature gradients, screw and gantry speeds, layer times, and nozzle designs were varied to minimize sharkskinning, warpage, and porosity of the final parts.

Print parameters including temperature gradients, screw and gantry speeds, layer times, and nozzle designs were varied to minimize sharkskinning, warpage, and porosity of the final parts. A strong effect of the nozzle size on the resulting porosity was observed, and consistent trends between decreasing porosity, increasing fiber alignment, and increasing mechanical performance were identified after printing with different nozzles, compression molding, and injection molding.IntroductionAlthough interest in both additive manufacturing (AM) and sustainable feedstocks for manufacturing has rapidly increased in recent years, the high cost of many bio-based polymeric materials in comparison to petroleum-based commodity plastics continues to restrict their use.

Large format additive manufacturing (LFAM) systems have also opened a variety of new applications areas for AM and new thermoplastic feedstocks as well.[1] Poly(lactic acid) (PLA) is a common feedstock for AM on small, desktop scales as well as LFAM systems and is currently one of the only completely bio-based thermoplastics available on a commercial scale.[2, 3] Discontinuous fibers or particles are often added to PLA and other resins to improve the mechanical properties of additively manufacture parts, add specific functionalities, or enable printability on larger scales.[4] The addition of natural fibers as opposed to synthetic to PLA also allows the composite to remain 100% bio-based and industrially biodegradable.

Wood flour (WF) in particular is a popular filler for AM feedstock due to its low cost, and PLA/WF can be readily purchased as filament for AM from such retailers as Amazon.[5, 6] The Oak Ridge National Laboratory, along with other research institutions, have conducted numerous LFAM studies using PLA filled with 20 wt.% of WF, and most of the commercially available PLA/WF filament is2113available at a filler loading of less than 20 wt.%.[7] However, these composites remain more expensive than commodity polymers such as polypropylene or polyethylene, but their performance (both mechanical and thermal) restricts their use in more demanding, higher value applications in which carbon fiber-reinforced composites are typically employed.[8] As such, this study aims to develop a lower-cost, bio-based composite feedstock for LFAM using higher loadings of natural fibers or particles and a small inclusion of additives for viscosity modification while retaining the mechanical performance of the composite.

Various loadings of WF in PLA were first investigated to determine a qualitative maximum limit of filler. A second type of natural filler, bleached Kraft pulp fibers, was also investigated for its improved reinforcing potential in comparison to WF. Finally, different viscosity modifiers were considered in the highly loaded composite to enable processability of the final composite. An optimal formulation was then scaled up to produce a large batch of feedstock for LFAM. Samples were prepared using the resulting composite by additively manufacturing with different processing conditions as well as compression and injection molding. Finally, the effects of processing methods on the performance of the final composite were compared.Materials and MethodsMaterialsPoly (L-Lactide), herein referred to as PLA (Ingeo 4043D), was purchased from NatureWorks, LLC. Wood flour (WF, 100 mesh, softwood pine) was acquired from American Wood Fiber.

Wood flour (WF, 100 mesh, softwood pine) was acquired from American Wood Fiber. Bleached Kraft pulp fibers (CreaTech TC750), herein referred to as Creafill, was acquired from Creafill Fibers Corp. Polyethylene glycol (PEG, 8000 g/mol) and poly(methyl acrylate-co-ethylene-co-glycidyl methacrylate) (P(MA-co-E-co-GMA)) were acquired from Sigma Aldrich. Wax flakes consisting of polyol fatty acids (Licolub WE 60FL) were acquired from Clariant AG.Sample preparation (bench scale)All materials were dried at 60 ℃ in a convection oven for a minimum of 4 hours before processing to remove moisture. Bench-scale samples were compounded using a melt mixer (Intelli-Torque Plasticorder half-size mixer, C. W. Brabender Instruments Inc.). PLA was first mixed at 180 ℃ for two minutes at 60 RPM, after which the desired additive (for some samples, PEG, PMA-co-E-co-GMA, or wax) was added and allowed to mix for another 3 minutes.

PLA was first mixed at 180 ℃ for two minutes at 60 RPM, after which the desired additive (for some samples, PEG, PMA-co-E-co-GMA, or wax) was added and allowed to mix for another 3 minutes. The desired amount of WF and/or Creafill was then added and allowed to mix for a further 5 minutes. The content of WF was first varied from 20 to 50 wt.% in the PLA without additives. Ratios of Creafill and WF at 100/0, 75/25, 50/50, 25/75 (Creafill/WF) were then compounded in PLA at 40 wt.% total fiber content. Finally, Creafill/WF at a 75/25 ratio was compounded in PLA at 40 wt.% total fiber with PEG, P(MA-co-E-co-GMA), and wax additives at 1 wt.% (separately).After compounding, each sample was compression molded into 1 mm films at 180 ℃ for 5 minutes using a hydraulic press (Carver). The films were then cut into approximately 3 mm strips and compression molded into uniform bars at 180 ℃ for 5 minutes following ASTM D4703.

The films were then cut into approximately 3 mm strips and compression molded into uniform bars at 180 ℃ for 5 minutes following ASTM D4703. Bars were routed into ASTM D638 Type V dogbones using a router (Tensilkut). Neat PLA pellets were also compression molded to Type V dogbones using the same procedure.2114Characterization (bench scale)The torque during melt mixing of select samples was recorded as a function of time by the Brabender for a mixing period of approximately 15 minutes. The tensile performance of each sample after compression molding was measured using servo-hydraulic equipment (MTS). Samples were extended at 1 mm/min (original gauge length of 7.62 mm, 13% strain/min) until fracture using a 50 N load cell.

Samples were extended at 1 mm/min (original gauge length of 7.62 mm, 13% strain/min) until fracture using a 50 N load cell. Five specimens of each sample were tested.Sample preparation (pilot scale)A target formulation was selected based on bench scale measurements, and a large batch of composite material (~100 lbs) was prepared using a corotating twin screw extruder (ZSE 27MAXX, Leistritz) with 12 heating zones. Temperatures along the barrel of the extruder were programmed from 145 ℃ at the inlet to 185 ℃ at the die, with a maximum temperature along the barrel of 190 ℃. The screw speed was set to 110 RPM. Six strands of composite were extruded and pelletized.Pellets compounded on the twin screw were compression molded following the procedure described above to produce bars according to ASTM D4703, some of which were routed to ASTM D638 Type V dogbones.

Pellets were also injection molded using a 35-ton machine (BOY 35 E VV) with 5 heating zones at temperatures ranging from 170-195 ℃. The pellets were plasticized at 300 mm/s to create a 30 mm shot, which was injected into an ASTM D638 Type I dogbone mold using 17 MPa of pressure.Hexagonal samples were printed using a Big Area Additive Manufacturing (BAAM, Cincinnati Inc) thermoplastic extruder. The BAAM consists of a single screw extruder mounted on a gantry system above a heated bed. Composite pellets were dried at 60 ℃ for 4 hours prior to printing. A polycarbonate build sheet was held in place on the bed with a vacuum system and heated to 60 ℃. The BAAM system consists of 5 heating zones which ranged from 155 to 220 ℃ (feeding zone to nozzle). One to three layers of neat PLA were printed initially to aid in adhesion of the composite to the print bed, after which the feedstock was switched to the composite pellets.

One to three layers of neat PLA were printed initially to aid in adhesion of the composite to the print bed, after which the feedstock was switched to the composite pellets. Three samples were printed using nozzles with diameters of 0.2” (5.08 mm), 0.3” (7.62 mm), and 0.35” (8.89 mm), respectively. The hexagonal samples were 10” on each side and were printed to a height of at least 8”. The screw speed was set to 30 RPM for each print, and the gantry speed was varied to improve the quality of the extrudate. The layer time during each print was approximately 1 minute.Characterization (pilot scale)Differential scanning calorimetry (DSC) measurements were performed on composite pellets and neat PLA using a TA Instruments apparatus. Samples were first equilibrated at 35 ℃, after which they were heated from room temperature to 200 ℃ at 10 ℃/min. samples were then cooled to 35 ℃ and heated again to 200 ℃ at 10 ℃/min.

Samples were first equilibrated at 35 ℃, after which they were heated from room temperature to 200 ℃ at 10 ℃/min. samples were then cooled to 35 ℃ and heated again to 200 ℃ at 10 ℃/min. The degree of crystallinity of both samples was calculated using equation 1,𝑋𝑐 =∆𝐻𝑚−∆𝐻𝑐 0∙𝜑∆𝐻𝑓(1)2115in which ΔHm and ΔHc denote the melting and cold crystallization enthalpies, respectively, of each 0 denotes the melting enthalpy of a 100% crystalline PLA sample, and φ denotes the material, ΔHf 0 for PLA was taken as 93 J/g.[9] The melting and cold mass fraction of PLA. The value of ΔHf crystallization enthalpies from the second heating scan were used to calculate the crystallinity of each material.The thermal degradation of the composite pellets and neat PLA was measured using a TA Instruments thermogravimetric analyzer (TGA). Samples were heated from room temperature to 500 ℃ at 10 ℃/min under argon flow. Measurements were conducted twice for each sample.

Samples were heated from room temperature to 500 ℃ at 10 ℃/min under argon flow. Measurements were conducted twice for each sample. The printed hexagons were machined to ASTM D638 Type I dogbones with their axes in the print direction (“x-direction”). The tensile properties of printed and injection molded (IM) samples were tested using a 2 kN load cell at 5 mm/min, while compression molded (CM) samples were tested using a 5 N load cell at 1 mm/min. Five samples of each type were measured.The viscoelastic response of AM, CM, and IM samples was measuring using dynamic mechanical analysis (DMA, TA Instruments). Samples were clamped with a dual cantilever clamp, and temperature sweeps were performed at a constant frequency of 1 Hz and amplitude within the sample’s linear viscoelastic envelope. Each sample was heated from 30 to 120 ℃ at 3 ℃/min.

Each sample was heated from 30 to 120 ℃ at 3 ℃/min. Measurements were conducted twice for each sample type.The internal structure and porosity of each sample was analyzed using x-ray computed tomography (xCT, Zeiss Metrotom) and VGEasyPore software. Samples were also cut from AM, CM, and IM specimens and potted in cylindrical molds with epoxy. After curing overnight, the samples in epoxy were polished to a mirror finish on a rotating autopolisher (Allied MultiPrep). The samples were then imaged using an optical microscope (Keyence), after which they were sputtered with platinum/palladium and imaged using a scanning electron microscope (SEM, Zeiss Merlin field emission) in secondary electron mode at an accelerating voltage of 1 kV.ResultsMaterial formulations were first iterated on a small bench scale to find an optimal material for scaling up.

PLA with 20 wt.% WF is already a very common AM feedstock, both on large scale system and small desktop printers.[10] In this study, we began formulation work with WF to determine the maximum possible loading in PLA. We then introduced Creafill fibers, which possess a much higher aspect ratio and have better reinforcing potential than WF, in mixtures with the WF. Finally, we tested various additives to the system to improve the melt flow and processability while maintaining as much mechanical performance as possible. As the primary goal of this work was to develop a composite with a high loading of natural fibers that remained processible via AM, the viscoelastic behavior is extremely important in addition to the performance of the resulting material.

It was found that samples were unable to be tested using conventional small angle oscillatory rheology, as the high solid content restricted the flow of the material between parallel plates and prohibited relaxation of the normal force on the instrument, meaning the material was unable to reach equilibrium before making a measurement. As such, the torque exerted by the motor of the melt mixing equipment was monitored. As each mixing procedure used identical temperatures and screw speeds, it was inferred that at higher torque2116reading was indicative of a higher shear viscosity for qualitative comparison of various formulations.Bench scaleSamples were first prepared using PLA and WF with 20, 30, 40, and 50% fiber by weight. The samples with 50 wt.% WF were found to be too brittle to successfully remove from the rectangular bar mold after compression molding. As such, all further samples were prepared with 40% fiber by weight.

As such, all further samples were prepared with 40% fiber by weight. Creafill was then added to the composites at varying ratios of Creafill/WF while maintaining an overall fiber loading of 40 wt.%. The torque during mixing of each sample was monitored on the Brabender instrument, and it was found that increasing the fraction of Creafill directly correlated to increasing torque in the equipment, which can be seen in Figure 1.Figure 1. Equilibrium torque (after 15 min) measurements of PLA with 20 and 40 wt.% WF and/or CreafillPLA with 40 wt.% WF and 40 wt.% Creafill was produced along with 3 mixtures of Creafill and WF at ratios of 25/75, 50/50, and 75/25. Each formulation was compression molded to produce dogbone samples for tensile testing, the results of which are listed in Figure 2.Figure 2.

Each formulation was compression molded to produce dogbone samples for tensile testing, the results of which are listed in Figure 2.Figure 2. Tensile testing results of PLA with 40 wt.% WF and/or Creafill (“/” refers to neat PLA with no filler)It is clear from the results in Figure 2 that Creafill is a better reinforcing fiber than WF in terms of tensile performance. Interestingly, the tensile strength and modulus both decreased2117slightly upon initial addition of Creafill (25/75 Creafill/WF sample versus 100% WF), but both increased at higher fractions of Creafill in the system. As WF particles have a much lower aspect ratio than Creafill fibers, they can likely flow more easily within a PLA melt in response to a given shear force, while the high aspect ratios of the Creafill fibers could lead to entanglement, increase tortuosity within the melt, and present an overall higher resistance to flow.

Based on the above results, a mixture of Creafill and WF at a ratio of 75/25 was chosen for future samples in order to balance the mechanical benefits of Creafill with the processability benefits of WF.Additives were then selected that had potential to improve the flow behavior of the highly loaded composite while maintaining as much of its mechanical performance as possible. PEG, P(MA-co-E-co-GMA), and a bio-based wax were then melt mixed with PLA and 40 wt.% Creafill/WF (75/25) at 1 wt.% of the additive (59% PLA, 40% fiber, 1% additive by weight).[11, 12] As with previous samples, the torque during mixing was monitored, and the tensile properties of each sample were tested, which are shown in Figures 3 and 4, respectively.Figure 3. Equilibrium torque measurements of PLA with 20 wt.% WF and 40 wt.% Creafill/WF (75/25) with additives (“/” refers to sample with no additives)Figure 4.

Equilibrium torque measurements of PLA with 20 wt.% WF and 40 wt.% Creafill/WF (75/25) with additives (“/” refers to sample with no additives)Figure 4. Tensile testing of PLA with 40 wt.% Creafill/WF (75/25) with additives (“/” refers to sample with no additives)The equilibrium torque of PLA with 20 wt.% WF is given in Figure 3 for reference, as PLA/20 wt.% WF is a common feedstock used in large- and small-scale AM, so its flow behavior was used as a qualitative benchmark in developing a printable highly loaded composite. As can be seen in Figure 3, the addition of wax to the highly loaded mixture significantly decreased its2118equilibrium torque, nearly to that exhibited by PLA/20 wt.% WF. Additionally in Figure 4, the addition of wax seems to slightly improve the tensile performance of the composite, although the values measured for tensile strength and modulus were within error of that of the unmodified composite.

It was thus decided that PLA with 40 wt.% Creafill/WF (75/25) and 1 wt.% wax would be scaled up for AM trials. Nearly 100 lbs of feedstock was compounded on a twin screw extruder after adjusting the extruder temperatures, pellet and fiber feedrates, and screw speed to maximize extrudate quality and throughput while maintaining the torque on the extruder motor and melt pressure within acceptable ranges (as dictated by the equipment manufacturer).Pellets compounded on the twin screw extruder were tested for their thermal behavior and stability using DSC and TGA, the results of which are shown in Figure 5. The thermal stability of Creafill and WF was also analyzed by TGA, included in Figure 5.Figure 5.

The thermal stability of Creafill and WF was also analyzed by TGA, included in Figure 5.Figure 5. DSC and TGA traces of neat PLA and PLA with 40 wt.% Creafill/WF and 1 wt.% wax and TGA traces of Creafill and WFAs can be seen in the first plot of Figure 5, the incorporation of fibers and wax additives to the PLA matrix significantly increases its cold crystallization and melting enthalpies. The degree of crystallinity within the composite sample was found to be 1.69%, compared to the almost negligible crystallinity within the neat PLA of 0.43%. The glass transition temperature was largely unaffected by the incorporation of fibers and additives to the system. The thermal stability of the material, however, was significantly affected by the addition of fibers and wax. As shown in the second plot of Figure 5, the onset of decomposition in neat PLA was at 337 ℃, while decomposition in the composite sample began at 292 ℃.

As shown in the second plot of Figure 5, the onset of decomposition in neat PLA was at 337 ℃, while decomposition in the composite sample began at 292 ℃. As can be seen in the last plot of Figure 5, Creafill is more thermally stable than WF, with an onset of degradation at 310 ℃ compared to2119WF’s 281 ℃. The lower thermal stability of the fibers in the PLA composite undoubtedly were responsible for its decrease in thermal stability in comparison to neat PLA.Composite pellets were then used to additively manufacture samples using the BAAM at Oak Ridge National Laboratory. The first set of test prints were conducted using a 0.35” diameter nozzle.

The first set of test prints were conducted using a 0.35” diameter nozzle. The initial temperature profile and screw speed were selected based on previous work on the same system with PLA/20 wt.% WF.[10] The torque on the BAAM motor and pressure at the nozzle were observed to be close to the upper limits of the system, so the temperature profile and screw speed were adjusted until a steady, consistent flow could be achieved. Severe sharkskinning was observed initially, which is shown in Figure 6.Figure 6. Initial AM samples exhibiting sharkskinning and inconsistent flowThe rough appearance of the extruded beads was corrected by iteratively lowering the gantry speed while keeping the screw speed constant (i.e., extruding more material per unit length of printed material). The composite material exhibited very poor adhesion with the polycarbonate build sheet which led to shifting during printing.

The composite material exhibited very poor adhesion with the polycarbonate build sheet which led to shifting during printing. As such, layers of neat PLA were first printed by hand-feeding PLA pellets into the extruder hopper before filling the hopper with composite pellets. The neat PLA layers can be observed at the base of the left sample in the second image of Figure 6. After parameter optimization, hexagonal samples were printed with three different nozzles with diameters of 0.35”, 0.3”, and 0.2”. An in-process image is shown in Figure 7 along with finished samples.Figure 7: AM process and printed hexagons2120The start/stop region of the hexagons was kept at the same point throughout the print.

The material flow became very inconsistent in the start/stop region, likely due to both the composite’s high viscosity and the internal settings of the BAAM extruder (gantry and screw are programmed to slow down at a set rate prior to abrupt directional changes/sharp corners), which could likely by overcome with further experimentation. Regardless, the areas affected by inconsistent flow in the start/stop region were discarded from further analysis.The three hexagons were cut and machined to produce dogbone specimens for tensile testing. Composite pellets were also compression and injection molded to produce tensile dogbones for comparison. Additionally, each sample (3 AM, 1 CM, 1 IM) was analyzed using xCT to image its internal structure and determine its overall volumetric porosity. Images from xCT of each sample are shown in Figure 8, in which black portions within the sample indicate porosity.Figure 8.

Images from xCT of each sample are shown in Figure 8, in which black portions within the sample indicate porosity.Figure 8. xCT images showing internal structure of CM, IM, and AM samplesIt can be clearly seen in Figure 8 that printing introduces a great deal of porosity into the material, while the CM and IM samples are seemingly free of porosity. The volumetric porosity of each sample was quantified and is shown in Figure 9, along with the measured tensile strength of each sample in its print direction.2121Figure 9. Tensile strength and porosity of AM, IM (“inj”), and CM (“comp”) samplesFigure 9 exhibits a clear trend of increasing tensile strength with decreasing porosity. The porosity of samples printed with the 0.35”, 0.3”, and 0.2” nozzles was found to be 50%, 41%, and 22%, respectively. No porosity was detected in the IM or CM samples.

The porosity of samples printed with the 0.35”, 0.3”, and 0.2” nozzles was found to be 50%, 41%, and 22%, respectively. No porosity was detected in the IM or CM samples. The decrease in porosity with decrease in nozzle size is likely due to the increase in pressure exerted by the smaller nozzles, leading to more compaction or consolidation of the material. Although the IM and CM samples were both nonporous, the tensile strength of the IM sample far exceeded that of the CM sample. This can be attributed to fiber alignment due to the high shear flow within the mold during injection molding.[13] The higher shear exerted by smaller nozzles during AM also likely caused fiber alignment and contributed to the improved tensile performance of the composites, although the positive effects of reduced porosity and fiber alignment are difficult to decouple.[14]Optical and SEM images of the polished sample surfaces are shown in Figure 10.

In the IM and AM samples, the flow direction (flow within the mold or print direction) of the composite during processing is labeled. Fibers can be seen in the higher magnification optical images (appearing white in the images) and SEM images, and there is clear alignment of fibers in each AM and IM sample with the direction of flow. Fibers in the CM sample, however, show no clear directionality. Additionally, macroscale porosity can be seen in the AM samples, labeled with blue arrows in Figure 10. Finally, the different layers of the print in the AM samples can be seen in the low magnification images, reflecting areas of high and low porosity as seen in the xCT images in Figure 8.2122Figure 10. Optical and SEM images of polished surfaces of CM, IM, and AM samples printed with different nozzle sizes.

Optical and SEM images of polished surfaces of CM, IM, and AM samples printed with different nozzle sizes. The flow direction on IM and AM samples is labeled with black arrows, porosity is identified by blue arrows, and fibers are identified by white arrows.ConclusionsA new bio-based feedstock was developed in this work, which sought to reduce the cost of bio-based AM feedstock by increasing the loading of low-cost natural fibers. The PLA used in this work was purchased from NatureWorks at approximately $1.60/lb, while the wood flour and Creafill were obtained at $0.65 and $0.85/lb, respectively. Based on these prices, the typical cost of PLA with 20 wt.% WF is roughly $1.41/lb, while the newly developed feedstock could be obtained at roughly $1.28/lb. Additionally, the embodied energy of the feedstock was also2123decreased by increasing the natural fiber content.

Additionally, the embodied energy of the feedstock was also2123decreased by increasing the natural fiber content. According to the Eco-Invent database, the embodied energy of PLA is 71.1 MJ/kg, while the embodied energy of bleached Kraft pulp is 46.96 MJ/kg, and the embodied energy of WF can be assumed to be 0, as it is considered a waste material. As such, the embodied energy of PLA with 20% WF is approximately 58.9 MJ/kg, while the embodied energy of the new material in this work is 56.7 MJ/kg. These cost and energy savings could be significant in large or high-volume projects involving tens of thousands of pounds of feedstock. However, these estimates are based on a weighted average of the cost and embodied energies of each constituent material and do not include considerations for packaging, transportation, or drying of materials.

The cost and energy of the composite developed in this work could vary greatly depending on the locations of the wood sources, composite production, and processing. Different sources of natural fibers could be considered for similar composites that are local to the composite production and processing site(s), in which case the formulation, processing conditions, and material performance would need to be re-evaluated.Even more significant was the effect of processing method on the performance of the material, which was shown to be related to both fiber alignment and porosity. The newly developed feedstock, consisting of PLA with 40 wt.% Creafill and WF at a 75/25 ratio, respectively, and 1 wt.% wax, was compounded on a pilot scale system and used to prepare samples using compression and injection molding as well as additive manufacturing on a large format system with three different nozzle sizes.

The highest tensile performance was observed in the injection molded sample, followed by the compression molded sample and the three additively manufactured samples in order of increasing nozzle size, meaning the AM samples with the smallest nozzle size showed the best performance. Additive manufacturing with pellet fed systems is known to result in porosity, which can vary based on numerous factors including the size and shape of the pellets, design of the feed system, fiber loading, fiber length, screw/barrel/nozzle design, and processing parameters.[15] It was theorized that porosity induced by AM was responsible for the inferior tensile properties as compared to IM and CM, and the fiber alignment induced by IM led to its higher performance in comparison to the CM samples.

The porosity of each sample was visualized using x-ray computed tomography, which proved that the IM and CM samples were virtually free of pores, while the porosity in the AM samples was positively correlated to the nozzle size. Fiber alignment was visualized by potting samples in epoxy and polishing them to a mirror-like finish. The polished samples were then imaged using optical microscopy and SEM. Again, no pores were visible in the IM and CM samples, while porosity was clearly visible in the AM samples. Additionally, alignment of the fibers could be seen in the IM and AM samples, while the fiber orientation appeared random in the CM sample.Processing methods inducing higher shear and resulting in lower sample porosity resulted in the best mechanical performance of the material developed in this study. To produce materials via AM that could compete with the performance of those produced via IM or CM, the porosity of the material must be reduced.

To produce materials via AM that could compete with the performance of those produced via IM or CM, the porosity of the material must be reduced. Reducing the nozzle size was shown to decrease part porosity, but reducing the nozzle size results in smaller layers and thinner walls, necessitating more layer and possibly more walls to than printing the same part with a larger nozzle, which can drastically increase print time. However, there are other modifications that could be made to an LFAM system to increase the printed parts’ performance. The design of the screw could be changed to impart more shear on the material and alter its compression ratio. Additionally, a vacuum port or vent2124could be installed on the extruder to remove volatiles trapped in the barrel during extrusion. This is a very common practice during compounding with twin screw extruders and is known to improve the quality and melt strength of the extrudate.

This is a very common practice during compounding with twin screw extruders and is known to improve the quality and melt strength of the extrudate. These alterations, among others, would likely result in AM parts with low porosity and highly aligned fibers with mechanical performance rivaling that of injection molded materials.Acknowledgements This manuscript has been authored by UT-Battelle, LLC, under contract DE-AC05-00OR22725 with the US Department of Energy (DOE). The authors acknowledge the support from the DOE Office of Energy Efficiency and Renewable Energy, Advanced Materials and Manufacturing Technologies Office, and the Oak Ridge National Laboratory/University of Maine Hub & Spoke Program team. Microscopy studies were completed at the Center for Nanophase Materials Sciences, a DOE Office of Science User Facility.

Microscopy studies were completed at the Center for Nanophase Materials Sciences, a DOE Office of Science User Facility. The US government retains and the publisher, by accepting the article for publication, acknowledges that the US government retains a nonexclusive, paid-up, irrevocable, worldwide license to publish or reproduce the published form of this manuscript, or allow others to do so, for US government purposes. DOE will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (https://www.energy.gov/doe-public-access-plan).References [1]S. M. E. Sepasgozar, A. Shi, L. Yang, S. Shirowzhan, and D. J. Edwards, "Additive Manufacturing Applications for Industry 4.0: A Systematic Critical Review," Buildings, vol. 10, no. 12, 2020, doi: 10.3390/buildings10120231. H. Storz and K. Vorlop, "Bio-based plastics: status, challenges, and trends," Landbauforschung Völkenrode, vol. 63, pp.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDevelopment of Joint Manufacturing and In-Line Metrology System for the Patterning of 3D Holographic Structures in Roll-to-Roll ProcessesBarbara Groh*, Kwon Sang Lee*, Michael Cullinan*, Chih-Hao Chang*Walker Department of Mechanical Engineering, University of Texas at Austin, Austin, TX, 78712AbstractRoll-to-roll (R2R) fabrication at the micro and nano scales promises to increase manufacturing throughput and reduce unit cost while providing avenues for unique product applications. By exploring the potential of creating 3D structures with a single lithography step and being able to confirm success in-situ, existing multilayer patterning error can be mitigated, since 3D features would be created in one step rather than many.

This paper demonstrates steps being taken to combine a R2R 3D nanolithography tool and an atomic force microscopy (AFM)- based in-line metrology tool into a functional system for patterning precise 3D structures. An existing manufacturing system will be adapted to pattern complex structures with a flexible PDMS mask currently being proven on stationary substrates. Modifications to the AFM system will include a focus on imaging patterns with varying mechanical properties and tailoring the system to include gathering mechanical information as well as imaging. By mapping surface features, the AFM tool will identify surface imperfections and predict failure modes occurring within the 3D structure.IntroductionCurrent roll-to-roll (R2R) manufacturing uses flexible webs made of thin sheets of metal, glass, polymers, or biomaterials stretched across a series of rolls to enable high-area, high- throughput manufacturing of flexible products.

However, flexible substrates have inherent properties that create significant challenge to the production of precise microscale and nanoscale patterns, and further limit the ability to maintain multilayer alignment [1].The challenges to high precision in R2R manufacturing mainly arise from the instability of the web once it is stretched across the rolls and present themselves in the form of flutter, warping, twisting, slipping and stretching [2], [3]. These instabilities result in a substrate that is difficult to control and monitor at the nanoscale. Fortunately, these challenges have been widely studied and may be accounted for in single-layer patterning applications with precision down to tens of nanometers [4].Since single-layer patterning can be reliably used in R2R, forms of lithography that create usable products in a single exposure prove to be ideal for this manufacturing process.

A commonly used lithography process for this kind of manufacturing is nanoimprinting [4], [5]. Another form of near-field 3D lithography uses near-field interference to leverage optical interference in creating 3D nanostructures within a resist during a single exposure to light [6]. This manufacturing ability will significantly reduce the necessary instances of multilayer alignment and reduce the risk of misalignment ruining a product.The goal of this research is to design and prototype a R2R tool capable of leveraging this form of interference lithography while simultaneously providing in-line metrology to improve the1034wholistic control of the manufacturing process and maintain a high-quality final pattern. In-line metrology will help mitigate the risk of failed patterns by monitoring for errors induced by the web instabilities and serve as one of many inputs to the process control algorithm of the system.

By building a joint manufacturing and metrology production system, errors will be prevented from propagating across the entire roll and prevent large-scale manufacturing losses.Manufacturing ProcessThe proposed manufacturing tool will be composed of two webs that interact within the lithography module to create the desired pattern. A model of the proposed manufacturing system can be seen in Fig. 1. In the diagram, the green web represents the conformal nanostructured optical mask made of polydimethylsiloxane (PDMS) that acts as an interference lithography mask. This flexible mask will loop to create a continuous mask for the lithography process to occur. The red web in the diagram is the flexible substrate on which the final pattern will be made—and which will proceed beyond this module to development and metrology modules not shown in Fig. 1.

The red web in the diagram is the flexible substrate on which the final pattern will be made—and which will proceed beyond this module to development and metrology modules not shown in Fig. 1. This mask will be coated in a layer of photoresist thick enough for the full 3D geometry of the near- field interaction to cure the full 3D structure.Fig. 1 Diagram of proposed manufacturing systemCurrently, the system is creates samples in a batch-to-batch process and does not operate on a continuous R2R system. Instead, every sample is created on a silicon wafer on a sample-by- sample basis. However, the processes necessary to prove the potential success of the R2R setup have been accomplished. The first step of the process is to create a silicon mold of the mask using Lloyd’s mirror lithography and RIE processing and is pictured in Fig. 2.1035Fig.

The first step of the process is to create a silicon mold of the mask using Lloyd’s mirror lithography and RIE processing and is pictured in Fig. 2.1035Fig. 2 Silicon nanopattern moldThe silicon mold can then be used to create an inverse PDMS mask by replicating the pattern using soft lithography. The flexible PDMS is at higher risk of damage due to high volume of use and the inherent instability of soft, flexible materials, but the silicon mold allows for nanopattern mask replication at a much faster time-frame than traditional mask manufacturing techniques. This is because the mold can be used repeatedly to form large-area masks rather than using slower processes like laser writing or e-beam writing. An example of the PDMS mask pattern can be seen in Fig. 3.Fig.

An example of the PDMS mask pattern can be seen in Fig. 3.Fig. 3 Conformal nanostructured PDMS optical maskThe final resist structure is made as a result of the PDMS mask, which is the final step of the nanopattern creation process and the ultimate product of the proposed nanolithography module shown in Fig. 1. Currently, the PDMS mask is placed on top of a silicon wafer coated with photoresist and then exposed to an ultraviolet (UV) light source on an optical table, creating the near-field holographic structures within the photoresist. However, this process will eventually be integrated into the R2R-based process as the full lithography module is constructed. A sample result of the 3D pattern can be seen in Fig. 4 as seen from above in a scanning electron microscope (SEM). Defects and inconsistencies in the proof-level testing process, including spin-coating errors and dust particle contamination, have resulted in samples with high variation.

Defects and inconsistencies in the proof-level testing process, including spin-coating errors and dust particle contamination, have resulted in samples with high variation. However, as the process is improved and optimized for adaptation to R2R fabrication, the resulting patterns will significantly improve in quality.1036Fig. 4 Final image of 3D holographic lithography patternMetrologyIn the interest of working towards a manufacturing tool with self-correcting process control, the development of in-line metrology methods is critical to the success of the research. The assessment of the quality of the patterns takes place in two main branches, atomic force microscopy (AFM) metrology and scatterometry. AFM will target feature-level metrology and take place on a quasi-continuous basis over predetermined sections of the web. Larger-area metrology will be targeted by scatterometry to provide faster results on the quality of the entire web.

Larger-area metrology will be targeted by scatterometry to provide faster results on the quality of the entire web. Scatterometry will provide the added benefit of providing insight on the success of internal structures within the 3D pattern. By working in tandem, the larger-scale metrology can serve to indicate pattern inconsistencies while the feature-scale metrology can determine specific defects that arise from various process failures.AFM Metrology:In-line AFM metrology of a moving web is only achievable when the AFM tip can move in tandem with the web to maintain a constant position relative to the web. This relative stillness is critical for the relatively slow imagining process to have a sufficient amount of time with the same scan area to get a full image. This tool uses a nanopositioning stage consisting of double- parallel flexures driven by voice coil linear actuators to provide XZ motion for AFM tip position control [7].

This tool uses a nanopositioning stage consisting of double- parallel flexures driven by voice coil linear actuators to provide XZ motion for AFM tip position control [7]. A picture of the flexure-based motion system is shown in Fig. 5, where the flexure system is further pictured to suspend a gantry system on which a laser distance sensor and the AFM itself are mounted.1037Fig. 5 Double-parallelogram nanopositioning flexure stageThe AFM chip used in this setup is a “single-chip AFM” or sc-AFM (ICSPI Corp.). This chip is a MEMS device with microscale 3D motion driven by thermal actuators and cantilever deflection detected by piezoelectric sensors. An image of the MEMS device is shown in Fig. 6, where the device is mounted on a chip that is mounted to the full-scale machine.Fig. 6 MEMS sc-AFM [8]Testing of AFM imaging with moving scans is still underway, but static scans of samples made using the holographic interference lithography process can be seen in Fig. 7.

6 MEMS sc-AFM [8]Testing of AFM imaging with moving scans is still underway, but static scans of samples made using the holographic interference lithography process can be seen in Fig. 7. This figure shows an sc-AFM scan of a 60 mJ/cm2 exposed sample with an un-sharpened tip (radius > 100 nm) and demonstrates the device’s ability to detect defects including missing features and pattern collapse.1038Fig. 7 Static sc-AFM scan of 60 mJ/cm2 exposed sampleAFM scans are limited to areas under 100 um2 every 15-30 seconds and cannot be repositioned relative to the width of the web, and therefore provide only small-area sample to draw process control conclusions.

A larger-scale metrology method is necessary to establish the health of the whole pattern, and scans within the visible spectrum of light will work to provide that information.ScatterometryScatterometry will not be able to provide the detailed characterization of a specific area that AFM scans will, but will be able to provide continuous feedback across the entire width of the web. Additionally, comparisons to optical models will allow the system to collect information on the success of the internal structure of the pattern where AFM scans are limited to surface-level detection.Scatterometry measurements result in reflectance spectra that modulate in period and amplitude that is dependent on both wavelength of light used for detection as well as sample dosage, which can be seen in Fig. 8.

8. By monitoring the modulation phase and amplitude, it will be possible to determine whether the sample was under exposed or over exposed, which would have resulted in internal pattern defects that are reflected in the reflectance spectra.Fig. 8 Reflectance measurements of fabricated samples with 70, 90, and 110 mJ/cm2 exposure dose versus wavelength.1039It is important to note that the reflectance spectra demonstrated in Fig. 8 were conducted using a slightly different 3D pattern sample than the samples generated in the manufacturing section. However, the samples were able to serve as proof of the success of the process and the unique reflectance spectra of the PDMS-masked patterns will be taken as the manufacturing process continues to evolve.Future Work: Joint Design ProposalWhile the manufacturing and metrology processes currently remain separate, plans to adapt these tools to function as a whole-scale system are underway. Fig.

Fig. 9 outlines a rudimentary system design with all major manufacturing and metrology components. These include the lithography module that uses the manufacturing steps outlined in the manufacturing section, including the continuous PDMS mask and the pre-coated web. This system does not currently include a photoresist application step, as a separate R2R slot-die coating system will coat the substrate roll before it enters the proposed tool.Fig. 9 Diagram of combined manufacturing and metrology systemAfter exposure, the web must enter a developer step to remove the uncured resist. This is immediately followed by a rinsing step to ensure that excess developer and resist molecules do not interfere with the metrology modules and final pattern.

This is immediately followed by a rinsing step to ensure that excess developer and resist molecules do not interfere with the metrology modules and final pattern. After rinsing, the web will encounter the metrology modules, which consist of the AFM and scatterometry for multiscale metrology.Not pictured is the process control algorithm that will take the outputs of the metrology modules, as well as countless other sensors system-wide, to determine the success of the pattern and modify any system control inputs to optimize the manufacturing process and improve the quality of the pattern.This level of process control is only possible with in-line metrology in a system as proposed. Out-of-line metrology would require the patterning of an entire roll before transfer to a separate metrology system, or cutting of the web that may interfere with the roll motion of any subsequent manufacturing.

Such delay in metrology could result in significant losses since pattern defects from imperfect process parameters would not be detected until many meters of pattern1040have been processed. By providing real-time feedback of pattern quality, such losses can be significantly prevented.ConclusionsWhile still in the early stages of development, individual process success demonstrates the feasibility of a 3D holographic lithography tool capable of multiscale metrology for process control optimization. Such a tool would greatly enhance the feasibility flexible-substrate manufacturing in industry by allowing for continuous and reliable 3D structure fabrication.References[1] line, Tip Based Nanometrology of Patterned Materials and Devices”.L. G. Connolly, A. Cheng, J. Garcia, and M. A. Cullinan, “A Roll-to-Roll System for In-[2] Storage Devices,” Advance Energy Materials, vol. 6, no. 23, Jul. 2016.J. Deng and L.

VELOPMENT OF LASER CLADDING PROCEDURE THROUGH EXPERIMENT AND ANALYSIS USING POWDER BLOWN DIRECTED ENERGY DEPOSITIOND. Garcia1,2*, K. I. Watanabe1,2, L. Marquez1,2, E. Arrieta1,2, R. Wicker1,2, and F. Medina1,21Aerospace and Mechanical Engineering, University of Texas at El Paso, El Paso, TX 79968, USA. 2 W.M. Keck Center for 3D Innovation, El Paso, TX, USA.Corresponding author: Email: dgarcia68@miners.utep.eduAbstractDirected Energy Deposition (DED) is one of the categories in Additive Manufacturing (AM) that has increased its popularity due to the technological advancements in recent years mainly with advancements in laser power, application of multi-materials, and capability to print greater dimensions. We are experimenting with a particular process in DED, cladding.

We are experimenting with a particular process in DED, cladding. For our experiment, we implemented a powdered feedstock (Inconel 718) which was assessed with distinct types of substrates in an effort to reduce the time spent on trial-and-error development of cladding parameters. We developed a procedure to determine a good clad interaction after an examination of the microstructure and interaction of single beads and a hatched area. The results demonstrate the ideal powers to be applied in the three substrates assessed with correlation to the dilution percentage where our target ranged from 10 to 30%.Keywords: Inconel 625, Inconel 718, Directed Energy Deposition (DED), Cladding, Additive Manufacturing (AM)IntroductionAdditive manufacturing (AM) is the process of producing three-dimensional objects which commonly builds one layer at a time.

This technology is aided by the application of 3D scanners or computer-aided design (CAD) which allows the system to manufacture complex geometric shapes [1]. In recent years, additive manufacturing has demonstrated rapid growth and interest among researchers due to its broad variety of materials, cost-effectiveness, and applications. Additive manufacturing can utilize a wide variety of materials, such as polymers, ceramics, metals, and composites. For these reasons, additive manufacturing has several applications in different industries, such as aerospace, transportation, medical, energy and consumer products. Furthermore, additive manufacturing can reduce lead times for wrought materials, significant cost and schedule reductions, printing components no longer built by the Original Equipment Manufacturers (OEMs), reduction of part by integration, etc [2].

Today, there are different additive manufacturing technologies, such as binder jetting, material extrusion, powder bed fusion, sheet lamination, and direct energy deposition.1107Scope of the StudyIn this research paper, our attention is focused on directed energy deposition. Specifically, in the cladding process between three varied materials which are: Inconel 718 as our feedstock, and Inconel 625, and Stainless-steel as the substrates. In a proper cladding process, it is critical to understand the proper interaction of the coating onto the substrate. Cladding is a directed energy deposition (DED) process in which metallic powder or wire feedstock is melted on metallic surfaces using a high-power laser to repair damaged surfaces or to enhance surface properties [3]. In this experiment our metal coating was Inconel 718, which is known for its high strength.

In this experiment our metal coating was Inconel 718, which is known for its high strength. Also, the process itself is one of the most time-consuming challenges; this involves finding the proper hatch width, powder feed rate, travel speed, deposit height, and understanding the interface and adhesion of the metal powder being applied onto the substrate. As mentioned, the purpose of cladding is to improve the properties of the material (yield strength, ultimate tensile strength, wear and corrosion resistance) to which the cladding is applied [4]. Being able to standardize the process of cladding by using powder blown directed energy deposition is significant. Some benefits that are achieved by this research not only include the optimization of the time-consuming process of adjusting the parameters, but also analyzing an acceptable and repeatable interaction and thermal characteristics.

Additionally, obtaining a good interaction would lead to enhanced properties in the substrate, moreover a proper restoration when repairing a geometry. This research would lead to a reliable and even cost- effective process.Materials and MethodsFor this experiment the machine utilized was an RPM Innovations 222XR Laser Deposition System. The machine itself operates in 5-axis and utilizes Laser Deposition Technology (LDT) defined as the process in which metal powder is injected into the focused beam of a high-power laser under tightly controlled atmospheric conditions to perform laser repair, laser cladding, and laser free-form manufacturing [5]. The machine’s gantry is capable of printing 2ft in the X-axis, 2ft in the Y-axis, 2ft in the Z-axis, and is equipped with a tilt table capable of 220⁰ of absolute tilt position, and a rotate table capable of continuous, absolute positioning.

The machine uses a 2-kilowatt Ytterbium fiber laser with a spot size ranging from 0.508 to 3.55 mm to melt powdered metal delivered by argon gas in a precision-controlled tool path. Oxygen levels were kept below 10 ppm to produce proper quality depositions during the experimentation process.Powder FeedstockInconel 718 powder feedstock with a size distribution of 45-150µm from AP&C (Boisbriand, Canada) was used for the study. The powder was produced by the Advanced Plasma Atomization (APA™) process where plasma torches melt and atomize metal wire feedstock sourced from 100% virgin melted material [6].Sample Identification and DocumentationA total of three experiments were studied for this research paper. The first experiment utilized Inconel 718 as the cladding material over Stainless-steel. The second and third portions of1108the experiment involved Inconel 718 on machined and as-built Inconel 625 plates.

The second and third portions of1108the experiment involved Inconel 718 on machined and as-built Inconel 625 plates. The spot size utilized over the three samples was 1.778 mm with a spacing between each clad of 7.62 mm for the Stainless-steel and 5.08 mm for the Inconel 625 substrates. Also, the mass flow rate used among the three experiments was 11 g/min. The parameters followed for these experiments are included in Table 1 and Table 2.Table 1. Inconel 718 on Stainless-steel (As-built specimen parameters)NameTest 1 Test 2 Test 3 Test 4 Test 5 Test 6 Test 7 Test 8Power (W)400 600 800 1000 1070 1200 1400 1600Speed (mm/s)5.08 7.58 10.12 12.66 13.55 15.20 17.74 20.28Energy Dens. (J/mm) 78.74 79.18 79.07 79.00 78.99 78.96 78.93 78.90Feed Rate (RPM) 5 5 5 5 5 5 5 5Table 2.

(J/mm) 78.74 79.18 79.07 79.00 78.99 78.96 78.93 78.90Feed Rate (RPM) 5 5 5 5 5 5 5 5Table 2. Inconel 718 on Inconel 625 (As-built and Machined)NameTest 1 Test 2 Test 3 Test 4 Test 5 Test 6 Test 7 Test 8 Hatch 1 Hatch 2 Hatch 3Power (W)400 600 800 1000 1070 1200 1400 1600 800 1000 1200Speed (mm/s)5.08 7.58 10.12 12.66 13.55 15.20 17.74 20.28 10.12 12.66 15.20Energy Dens. (J/mm) 78.74 79.18 79.07 79.00 78.99 78.96 78.93 78.90 79.07 79.00 78.96Feed Rate (RPM) 4.7 4.7 4.7 4.7 4.7 4.7 4.7 4.7 4.7 4.7 4.7Experiment SetupSamples for metallography observation were first mechanically grinded, then polished, and etched. The grinding and polishing procedures were conducted with an ATM SAPHIR 530 semi- automatic system. For test 1; Inconel 718 on Stainless-steel, Kalling's No. 2 (see Table 3) is used for Inconel 718 and Carpenter’s Etchant (see Table 5) for Stainless-steel 300 series.

For test 1; Inconel 718 on Stainless-steel, Kalling's No. 2 (see Table 3) is used for Inconel 718 and Carpenter’s Etchant (see Table 5) for Stainless-steel 300 series. Test 2; machined Inconel 718 on Inconel 625 and Test 3; as-built Inconel 718 on Inconel 625, Etchant 1 (see Table 4) is used.1109The grinding procedure started with a 320-grit coarse Si-C (Silicon carbide) grinding paper, followed by grits 600, 800, and 1200, each was spun at 200 RPM with a force of 22 N for 1 minute using water as a lubricant. The polishing procedure was conducted at a speed of 200 RPM and a force of 22 N for 1 minute, with each stage of the polishing process having a different disk for 6 μm, 3 μm, and 1 μm diamond suspension. Finally, the samples were polished for 30 seconds at 100 RPM with a force of 22 N using a TRICOTE polishing pad and a 0.05 μm Al₂O₃ (Alumina) suspension.

Finally, the samples were polished for 30 seconds at 100 RPM with a force of 22 N using a TRICOTE polishing pad and a 0.05 μm Al₂O₃ (Alumina) suspension. The samples were placed in an ultrasonic bath in between each polishing stage to eliminate cross-contamination.Kalling’s No. 2 Etchant consists of 5 g of Cupric chloride, 10 ml of hydrochloric acid, and 100 ml of Ethanol [7]. For the etching procedure in this investigation, the specimens were swabbed with the etching solution for a range of 30 to 60 seconds. Carpenter’s Etchant consists of 8.5 g of Ferric chloride, 2.4 g of Cupric chloride, 122 ml of Hydrochloric Acid, 6 ml of Nitric Acid, and 122 ml of Ethanol. The immersion of the sample was placed in the etchant for 5 seconds left out in air for another 5 seconds, then placed in DI water. Etchant 1 was used in electro etching consisting of 70 ml Phosphoric acid and 30 ml Deionized water at room temperature.

Etchant 1 was used in electro etching consisting of 70 ml Phosphoric acid and 30 ml Deionized water at room temperature. Using 5 volts for 12s- 2mins depending on the size of the sample.222XR RPM machine (DED) machine was used for the experiment setup for the cladding process. For test 1 the substrate utilized was a Stainless-steel plate. The plate was clamped into the bend of the machine by using two clamps holding the plate from both sides. The process took a total of 8 single bead tests replicated 3 times along the plate. Furthermore, the 2 last tests involved the cladding process of a hatch. An overlap of 33% was utilized for all the three hatches, each running at different power and speed. The same process was repeated for the machined plate.Table 3. Kalling's No. 2 Table 4. Etchant 1CuCl2Hydrochloric acid5g10 mlPhosphoric acid Deionized water70 ml30 mlEthanol100 mlTable 5.

Kalling's No. 2 Table 4. Etchant 1CuCl2Hydrochloric acid5g10 mlPhosphoric acid Deionized water70 ml30 mlEthanol100 mlTable 5. Carpenter's EtchantFeCl3 CuCl2 Hydrochloric acid Nitric acid Ethanol8.5 g 2.4 g 122 ml 6 ml 122 ml1110Results and DiscussionMicrostructure CharacterizationTests were conducted on eight different power and speed traits. ImageJ software was utilized for data analysis on the single beads and hatched areas. Ideally, our focus was on obtaining a dilution between 10 to 30% [8]. Dilution is described as the proportion of base material in the resultant weld metal, and for a single bead deposit, it is usually taken to be the ratio of the cross- sectional area of melted base material to the total cross-sectional area of the fusion zone [9]. Figure 1 depicts the formula utilized to calculate dilution percentage.Dilution =ABottom ATop + ABottomFigure 1.

Figure 1 depicts the formula utilized to calculate dilution percentage.Dilution =ABottom ATop + ABottomFigure 1. Dilution FormulaTest 1 - Inconel 718 on Stainless-SteelThe following single beads were obtained on the Stainless-Steel substrate. The powersused ranges from 400 W to 1600 W (see Figure 2).(a)(b)(c)(d)(e)(f)(g)(h)Figure 2. Single Beads on Stainless-steel: (a) 400, (b) 600, (c) 800, (d) 1000, (e) 1070, (f)1200, (g) 1400, and (h) 16001111400 W600 W800 W1000 W1070 W1200 W1400 W1600 WFigure 3. Microstructure of Single Beads on Stainless-steelFigure 3 portrays the profile of the single beads done on the Stainless-steel build plate. Additionally, on Figure 4 we were able to identify a decrease in the height of the samples as power was increased. At the same time, the width of each of the samples increased.Figure 4. Height & Width vs.

At the same time, the width of each of the samples increased.Figure 4. Height & Width vs. Power Graph on Stainless-steel1112After data was obtained, a graph (see Figure 5) was plotted to show the interaction of dilution (red dotted line) and the contact angle (blue dotted line) in contrast to the power used in each of the samples. As mentioned before, a proper dilution is targeted to be around 10 to 30% [8]. It was concluded that the ideal candidate for a proper cladding ranged from 800 to 1200 W (see area covered by blue rectangle) for the Stainless-steel substrate.Figure 5. Dilution vs. Power Graph for Test 1Test 2 - Machined Inconel 718 on Inconel 625For test number two, the same process was followed. Once the data was obtained, a graph was plotted to identify the proper power range according to the dilution percentage previously defined. Figure 6 shows the data obtained from the graph.

Once the data was obtained, a graph was plotted to identify the proper power range according to the dilution percentage previously defined. Figure 6 shows the data obtained from the graph. Interesting enough, unlike Stainless- steel it was found that the ideal candidates ranged at a lower power; 600 to 800 W (see area covered by blue rectangle). Additionally, when comparing Figure 4 and Figure 7, it is notable that the cladding interaction on the Inconel 625 substrate it is wider and shorter than the Stainless-steel substrate when ranging from 400 to 1200 W.1113Figure 6. Dilution vs. Power Graph for Test 2Figure 7. Height & Width vs. Power Graph on Machined Inconel 625Test 3 – As-built Inconel 718 on Inconel 625Unlike the machined Inconel substrate, the as-built specimen showed a lower range of power for a proper dilution.

Power Graph on Machined Inconel 625Test 3 – As-built Inconel 718 on Inconel 625Unlike the machined Inconel substrate, the as-built specimen showed a lower range of power for a proper dilution. As mentioned, in Figure 8 we can observe that the power ranged from 400 to 600 W (see area covered by blue rectangle), obtaining dilution percentages values of 16.52 to 25.46%.1114Figure 8. Dilution vs. Power Graph for Test 3Moreover, when comparing both substrates for the Inconel 625 (machined and as-builtsamples) we observed a distinctive pattern in its density. Figure 9 depicts the increase in porosity (see area covered by blue rectangle) in the as-built sample of Inconel 625. On the other hand, the machined samples presented less porosity. In essence, these results can be correlated to their manufacturing process. The additively built specimen was expected to have a lower density compared to the machined one as it didn’t have any proper post-processing.800 W800 WFigure 9.

The additively built specimen was expected to have a lower density compared to the machined one as it didn’t have any proper post-processing.800 W800 WFigure 9. Comparison of Inconel Machined (Left) and As-built Samples (Right)ConclusionParameters obtained both gantry speed and power) are intricately connected to chemical compatibility between the known substrate and the cladding material intended to be used. Also, a key factor commonly underestimated is surface finish. As already discussed, there was a clear quality and density improvement on the bond between the substrate and cladding material on the machined EBM Inconel 625 substrate. Parameters shown on the described research must be1115analyzed and confirm that optimization of dilution does not considerably affect sample’s microstructure.Further research on the interaction of the two materials is intended to be conducted.

For instance, one of the experiments that could be done is using Inconel 718 on an Inconel 625 cylinder with a vertical setup. The target of this experiment would involve building on top of the Inconel 625 bar with the purpose of assessing the interaction between the two materials by conducting a hardness test. Also, to understand the effect of the vertical building of the clads and how are the properties affected.1116References[1] What is Additive Manufacturing? definition, types and processes. TWI. (n.d.). Retrieved August 3, 2022, from https://www.twi-global.com/technical-knowledge/faqs/what-is- additive-manufacturing\[2] Gamon, A., Arrieta, E., Gradl, P. R., Katsarelis, C., Murr, L. E., Wicker, R. B., &amp; Medina, F. (2021, November 3). Microstructure and hardness comparison of as-built Inconel 625 alloy following various additive manufacturing processes. Results in Materials.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDevelopment of Multimaterial Additive Manufacturing Systems for Embedded ElectronicsCarson Vath, Kazi Md Masum Billah, Guhaprasanna ManogharanDepartment of Mechanical Engineering, Pennsylvania State UniversityDepartment of Mechanical Engineering, University of HoustonAbstractHybrid additive manufacturing (AM) to integrate discrete material systems and structures in a monolithic part is a growing research interest. The layer-by-layer deposition system of the AM build process allows users to integrate multiple materials using custom made tools added to the gantry systems. State of the art machine development efforts are highly focused on thermoplastic based material extrusion systems.

State of the art machine development efforts are highly focused on thermoplastic based material extrusion systems. Compared to the significantly matured thermoplastic material extrusion AM system, thermoset systems are not well positioned in the market due to the lack of integrated tooling. This research develops a wire deposition tool that is designed to embed wire in a photocurable thermoset. Thermoset is relatively stable and does not require heat for implanting wire. The proposed design method has a more rapid production rate as the extrusion process can be conducted without any interruptions. The final design was manufactured in PLA with a traditional FDM machine allowing multiple design iterations to be made quickly. The final design will be printed on an SLA machine for more accurate, robust parts.Introduction:Multimaterial additive manufacturing systems are one of the most highly sought after technologies in the manufacturing community.

Integration of the multiple materials in a monolithic structure within a single build is a unique characteristic of any AM system. Particularly, thermoplastic based material extrusion systems are ubiquitously developed from the desktop scale to large scale. Custom made tools are also developed and added to existing systems to incorporate multiple materials. Below we are providing brief yet to the points literature survey on the tool developments for the thermoplastic based material extrusion systems.1. Thermoplastic Deformation Based Wire EmbeddingThe two main methods for wire deposition are ultrasonic and thermal embedding. Both methods increase the energy of the substrate to melt the surface. Ultrasonic has been the most tested form of wire deposition. Tests in multi-axis conformal deposition have been successfully performed. [4] The ultrasonic deposition process rapidly vibrates a horn to create friction.

Tests in multi-axis conformal deposition have been successfully performed. [4] The ultrasonic deposition process rapidly vibrates a horn to create friction. [1] The friction heats the surface and allows for various materials to be embedded. Ultrasonic horns can deposit copper wire, conductive ink, and wire meshes. However, there are some drawbacks. While conductive ink is easier to implement, its electrical resistance is much higher than traditional copper wire. Additionally, if deposition of the ink is not consistent the trace could be interrupted leading to a poor connection or broken circuit. Due to the high frequency vibrations while1042embedding solid wire, wear on the horn and wire breakage were common. Thermal deposition works in a similar way. However, rather than using friction to heat the substrate, it uses a heating block to heat the wire which melts the surface. Because it is only heating the embedding material, only copper wire can be embedded.

Because it is only heating the embedding material, only copper wire can be embedded. [2] Due to the thermal conductivity of copper wire heat can travel down the wire and uproot previously embedded wire. That means this method has issues with sharp turns and retaining wire in the substrate. A difficulty that both methods face is the deformation of substrate surface. FDM relies on clean, smooth surfaces for printing. Rough or warped surfaces can lead to poor adhesion between layers causing cracks to form on the part. Rough surfaces can also cause clogging in the filament nozzle.2. Channel EmbeddingIn channel embedding a trench is first created during printing or machined with end milling. [1] Wire is then placed in the trench and retained by friction. A printing head will then deposit new material over the placed wire to secure it in place. Although this method is the simplest it has several issues.

A printing head will then deposit new material over the placed wire to secure it in place. Although this method is the simplest it has several issues. Most concerning is its struggle to retain embedded material both before and after the substrate is layered over it. Leading to poor traces and damaged circuits. If the channels are decreased in size, tool wear will increase.3. CoextrusionIn current coextrusion methods gallium liquid metals (LM) are fed through a nozzle and coated in SEBS elastomer. [3] The use of this method has not been tested in 3D printing and only the tool head was tested. Ensuring proper extrusion rate of each fluid was essential. If it was not perfect the liquid metal core would bead up and not become a continuous cylinder. Additionally, the syringe delivering the liquid metal had to stick out beyond the extruder or similar issues of beading would occur. At the end of the test flexible wires were created.

Additionally, the syringe delivering the liquid metal had to stick out beyond the extruder or similar issues of beading would occur. At the end of the test flexible wires were created. However, the liquid metal used was a factor of 100 times more expensive than traditional copper wire making it economically difficult to implement in large scale manufacturing.Proposed Method – Coextrusion of Solid Copper Wire with Photocurable ThermosetThe tool being created was based on the coextrusion method. Rather than using the LM a solid copper wire will be fed through the nozzle and in place of the SEBS a photocurable resin will be used. It is expected to overcome some existing challenges we surveyed in the thermoplastic systems. In our proposed design, there is no need for thermal energy to embed the wire. Therefore, it is expected there should be diminished warping of the print surface.

In our proposed design, there is no need for thermal energy to embed the wire. Therefore, it is expected there should be diminished warping of the print surface. Thermosets are known to be resistant to heat, which means if the wire is heated under electrical load, it will not shift. From the literature survey we learned that one of the largest issues faced by other methods is the accuracy of 90 degree turns [1,2]. While creating sharp corners using the proposed design the laying of wire will pause to wait for the curing of the thermoset. Once. It is cured the wire will have a new anchor point and will be able to be laid perpendicular to its old path.1043Design of the Tool – Machine Interfacing and ConsiderationsThe tool is designed to interface with the Hyrel Hydra 21. The Hydra has up to 5 head slots that are all attached to the XY carriage while printing. This means the tool's width must be constrained to 34 mm so that other tools can fit on either side.

The Hydra has up to 5 head slots that are all attached to the XY carriage while printing. This means the tool's width must be constrained to 34 mm so that other tools can fit on either side. Additionally, the tips of the heads being used for a single print must be on the same plane. If one is longer than the others it will crash into the bed or the part during printing causing damage to the print, the head, or the bed.Figure 1: Hydra 21 [5]Design of the Tool – Substrate ChannelsWith these constraints in mind a design was created. For coextrusion of the wire and gel there are two channels. The first channel, for the wire, is created by a syringe needle. The syringe for the original design is 23 gauge (0.337 mm) which was chosen to deposit 30 AWG wire (0.254 mm). This gives a loose enough fit while preventing the wire from buckling inside the needle. The syringe can be changed to deposit different gauges of wire. The second channel is set at a 20° angle from the wire channel.

The syringe can be changed to deposit different gauges of wire. The second channel is set at a 20° angle from the wire channel. This channel will carry the gel. The syringe will ensure that they stay separate until exiting the tip of the tool.Design of the Tool – Substrate Drive SystemThe wire will be driven much like a common FDM printer direct drive system. The gears for the direct drive were placed as close to the syringe entrance as possible to ensure the wire would not buckle. If it did buckle the wire would not pass into the syringe causing a clog to occur. The driving method for the gel is a syringe deposition tool already created by Hyrel. This tool head will connect to the wire deposition tool via a tube. Gcode will be created that will extrude the1044correct amount of gel for smooth deposition. The rate of deposition can be edited directly in the Hyrel software.

Gcode will be created that will extrude the1044correct amount of gel for smooth deposition. The rate of deposition can be edited directly in the Hyrel software. This makes it easy to run experiments at different deposition rates to determine which is the best for printing.Figure 2: Full Tool AssemblyParts List1. Main Body – 1x 2. Spacer – 1x 3. 28 mm x 28 mm stepper motor – 1x 4. Spool of wire – 1x 5. M2.5x45 screw - 4x 6. 7. Syringe needle – 1x 8. 3 mm diameter 20 mm length metal rod – 1x 9. Extruder bearings – 2x 10. Tensioner wheel – 1x 11. Tensioner door – 1x 12. Push-to-connect fitting – 1x 13. Hyrel software development kit (SDK) board – 1x 14. Hyrel board mount – 1x 1045.1 mm E3D nozzle – 1x15. Hyrel syringe deposition head – 1xAssembly DesignFor easy cleaning, wire insertion, and printing the design was split into two parts; the main body and a spacer. The main half of the body holds the channels for the syringe and gel as seen in figure 4.

The main half of the body holds the channels for the syringe and gel as seen in figure 4. Originally, the main body of the tool was split into two parts along the ZX plane of the syringe. This was so the syringe could be placed into one side of the body and then enclosed by the other half. However, concerns of gel leakage arose so it was determined that the syringe should be held by a single piece. There were also difficulties cleaning the internal structure, placing the syringe, and loading the wire with the single body design. Therefore, a part of the main body was cut away to allow easier access while maintaining proper dimensions for the bolts. These bolts are used to attach both main body parts to the stepper motor and act as a hinge for the tensioner door. The tensioner door houses an idler gear that adds pressure to the gear teeth driven by the stepper motor. The tension is created using a spring for a constant and easily adjustable force.

The tensioner door houses an idler gear that adds pressure to the gear teeth driven by the stepper motor. The tension is created using a spring for a constant and easily adjustable force. Once all parts are attached to the stepper motor the tool is attached to a Hyrel board interface. This interface connects directly to the Hydra 21 and allows the board to seat into the connector. Since all tool heads must be at the same height the tool is able to be adjusted in the z direction. The gel will be delivered via the Hyrel syringe deposition device as seen in figure 3. A hose will connect the syringe depositor to the tool body via a push-to-connect fitting.1046Figure 3: Hyrel Syringe Deposition Device [6]1047Figure 4: Tool Cross SectionElecronic Control SystemsTo interface the tool with the Hyrel system a Hyrel designed SDK board will be used. This board was designed to drive a typical FDM hot end but has the potential to drive a multitude of different systems.

This board was designed to drive a typical FDM hot end but has the potential to drive a multitude of different systems. The SDK board has an A4988 micro stepping motor driver which is a common motor driver used in many different 3D printers. This will allow the board to control the 28 mm stepper motor chosen for this application. The board also offers other connections such as 12 volts, 5 volts, PWM (Pulse Width Modulation) signals, RTD connections, and switches that can be controlled by sensor inputs. This will be vital for the future of the tool. To make it completely autonomous, a wire cutter will be designed, and other connections will be utilized.1048Figure 5: SDK Board [7]Software and Control MethodFor slicing parts Prusa Slicer will be used. Hyrel provides recipes for the software that allows Gcode to be generated for the Hydra 21 directly from Prusa Slicer. The slicer includes the ability to change heads during the print, which is vital for the tool's success.

The slicer includes the ability to change heads during the print, which is vital for the tool's success. This will allow the printer to extrude PLA from a traditional head then swap to the wire extrusion head and change back once the wire has been deposited. The software on the Hydra known as Repetrel will also be used to control the tool. Built into the Repetrel system is a cloning function. This function was intended to print multiple parts at the same time from multiple heads by choosing a master head and mimicking the extrusion on other heads. In this application the cloning function will be used to drive the syringe deposition tool. This will be accomplished by choosing the main body of the tool as the master tool and cloning those moves to the syringe. That means that whenever the wire stepper is activated the syringe stepper will also be activated.

That means that whenever the wire stepper is activated the syringe stepper will also be activated. To further control the gel, the extrusion rate can be overridden on the head software.1049Results and Future PlansThe above assembly was built to test fit and function. A UV curable solder mask was used as the thermoset during testing. Both extruded at the same time as planned, however, the mask was too viscous and did not hold its shape after extrusion. A thermoset with a higher viscosity is desired, but if it is too thick the fluid will be difficult to pump through the system. Therefore, the thermoset chosen should be shear thinning to allow for easier flow while under pressure. Additionally, since the main body was printed in PLA the push-to-connect fitting was not sealed properly and there was leakage. Finally, the needle was not secured with epoxy so some of the mask traveled up the needle channel.

Finally, the needle was not secured with epoxy so some of the mask traveled up the needle channel. Eventually, the syringe deposition will have to be replaced with a tool that has a larger reservoir as the current tool only holds 10 mL of liquid. A larger reservoir will prevent the need for refilling during longer prints. After testing, a larger gauge needle might be needed to allow for less friction as the wire passes through. Any defects in the wire increase frictional forces and prevent the wire going through smoothly.Figure 6: Mask Spreading Post CoextrusionReferences1050[1] Martinez, Nikki L., "Development of an Automated Ultrasonic Wire Embedding Process for use with Material Extrusion Additive Manufacturing and Rapid Electronics Fabrication" (2020). Open Access Theses & Dissertations. 3179.

velopment of nanocomposite powders for the SLS process to enhance mechanicalproperties Hyo Chan Kim1, Tsuyoshi Saotome1, H. Thomas Hahn1, Young Gil Bang2, Sung Woo Bae3, Dong Soo Kim3 1 Engineering IV, Mechanical & Aerospace Eng. Dept.,UCLA, L.A., CA90095, U.S.A. 2 Lion Chemtech., Co. Ltd., 41-5 Moonpyong-Dong, Daeduk-Gu, Daejeon, S. Korea 3 Korea Institutes of Machinery & Materials, 171 Jang-Dong, Yuseong-gu, Daejeon, S. KoreaReviewed, accepted August 27, 2007Abstract In an effort to fabricate prototypes with improved mechanical properties in the dual lasersintering process, functionalized graphite nanoplatelets were added to the PA-12 powder toproduce a nanocomposite powder. The PA-12 powder was chosen as the matrix polymer because it has features conducive to laser sintering such as relatively low melting temperatureand high mechanical properties. The GNPs were oxidized through a nitric acid treatment toimprove the interfacial bonding.

The GNPs were oxidized through a nitric acid treatment toimprove the interfacial bonding. The resulting nanocomposite powder was layered and sinteredby laser without any sign of agglomeration. Although the result is preliminary, it nevertheless shows the suitability of the nanocomposite powder for the laser sintering process.Key Words : Rapid prototyping, Nanocomposite, Graphite nanoplatelets, Polyamide-12,Dual Laser Sintering processIntroduction In recent years the product life cycle has become shorter than ever because of rapidly changingcustomer demands. Since it is almost impossible for conventional methods of productmanufacturing to meet these demands, much research is being done to reduce the time and cost of product development when diverse models and frequent design changes are indispensable.

Thesolid freeform fabrication (SFF) technology, also known as the rapid prototyping (RP) technologydeveloped since the late 1980s, has established itself as part of CAD/CAM and is expected tocope with the dynamic manufacturing environment. RP is an additive manufacturing process, in which a 3-D computer model is sliced and reassembled in real space layer-by-layer [1-2].There are many commercial SFF systems such as SLA (stereolithography apparatus), FDM(fused deposition modeling), SLS (selectivelaser sintering), LOM (laminated objectManufacturing), JP5 (JP System 5), 3DP (3D Printing) and some other processes [3]. However, all these processes are limited by the available materials although much research have been doneto make available better materials.Nanocomposites are composite materials where the reinforcement phase has at least onedimension in the range of 1-100 nm. There are many types of nanocomposites depending on thematrix materials and nanofillers used.

There are many types of nanocomposites depending on thematrix materials and nanofillers used. Integrating reinforcement particles in a polymer of choiceto form a composite material has been practiced for decades [4]. They have attracted muchinterest due to their special physicochemical properties which may be quite different from theirconstituent materials’. These composite materials can offer potentially useful and unusual combinations of mechanical, electrical, magnetic and optical properties that are otherwiseunachievable in conventional materials [5]. Due to many advantages of nanocomposites, the fieldof their application has been gradually expanding. The use of Nan composites in the SFFprocesses would be a timely development.A few researchers have tried to apply nanocomposites to the SFF processes. In 1993,Manthiram et al. [6] developed a nanocomposite to optimize selective laser sintering (SLS) andselective laser reactive sintering (SLRS).

In 1993,Manthiram et al. [6] developed a nanocomposite to optimize selective laser sintering (SLS) andselective laser reactive sintering (SLRS). They selected ceramic-ceramic and ceramic-metalnanocomposite systems in such a way that one nanosize component has a lower melting temperature than the other nanophase. The nanocomposite powders of Al2O3-CoO3, and Al2O3-NiO were synthesized by a sol-gel processing. In 2006, Ritzhaupt et al. [7] tailored the refractiveindex of polymethylmethacrylate (PMMA) using a UV-curable reactive resin and two differentalumina powders and one nanosized silica. In the same year, Kim et al. [8] developed a nanocomposite deposition system (NCDS) to fabricate 3D nanocomposite parts. The NCDS usesa photo-curable polymer resin as the matrix and various nanoparticles to form a compositeprototype.

The NCDS usesa photo-curable polymer resin as the matrix and various nanoparticles to form a compositeprototype. They used multi-walled carbon nanotubes (MWCNT) to improve electrical propertiesand hydroxyapatite particles to produce a biocompatible composite.In most previous works, nanocomposites used in the SFF process were fabricated by simplemixing of nanoparticles and polymer. However, this results in poor bond between the filler andthe polymer matrix, and introduces artificial defects, which can consequently be detrimental tothe mechanical properties of the final nanocomposite [9].In our work, a nanocomposite powder consisting of PA-12 (polyamide-12) powder andfunctionalized GNPs (graphite nanoplateletx) has been synthesized and evaluated for the duallaser sintering process. The PA-12 powder, which has good laser sintering characteristics such asrelatively low melting temperature and good mechanical properties, was chosen as the matrix polymer.

The PA-12 powder, which has good laser sintering characteristics such asrelatively low melting temperature and good mechanical properties, was chosen as the matrix polymer. To improve the interfacial bonding the GNPs were oxidized through a nitric acidtreatment. The resulting nanocomposite powder was used in the dual laser sintering process.Description of the dual laser sintering process Kim et al. [12] has developed the dual laser sintering process to increase the efficiency of lasersintering. The dual laser system which can scan two separate regions individually should beemployed in a SFF system capable of large size fabrication (e.g., 500x800x600 mm). The duallaser scanners improve the precision and efficiency of the SLS machine.As illustrated in Fig.

The duallaser scanners improve the precision and efficiency of the SLS machine.As illustrated in Fig. 1, the dual laser sintering system consists of a laminating module thatsupplies and transfers the powder, a heating module to preheat the powder, a nitrogen supplymodule to create a nitrogen atmosphere, a dual laser module that supplies the laser energy to alarge area, and a control module to control the entire system.Fig. 1 Schematic diagram of SFF systemThe powder is deposited layer by layer and a structure is constructed through sintering of the powder by the laser. Precise controls for powder deposition in the z axis and powder transfer inthe x axis by a roller are critical to the precision of a 3D structure. The layer deposition cyclestarts with the powder being moved up to the feeding room. The powder is then spread over thesupport by a roller in the building room.

The layer deposition cyclestarts with the powder being moved up to the feeding room. The powder is then spread over thesupport by a roller in the building room. The roller mechanism has several specially designed features that influence the roughness of the powder surface and the porosity of the interior. Thecritical parameters include the linear and rotational speeds of the roller, the feed ratio, the rollerroughness, and the deposition layer thickness.Laser sintering of a polyamide powder in the SLS process requires a build room temperature of 150 ±1 oC or higher to preheat the powder. The current machine utilizes a radiant heating system for preheating. The preheat temperature should be controlled precisely because any overshootmay cause curling or over-melting of the prototype.The nitrogen supply module is used to produce a nitrogen atmosphere in the work room and for the scanner lens.

The nitrogen atmosphere prevents the soot formation resulting from microexplosions that can occur during laser sintering and prevents the powder from exploding andsticking to the scanner lens. The nitrogen module maintains over 95 % nitrogen concentration inthe build room during operation.The laser head unit has been designed for the dual laser sintering. The laser module wasmanufactured using a 3-axis, dynamic focusing lens such that a 500×800×500 mm prototype could be fabricated using two laser beams. As shown in Fig.2, the laser module consists of a CO2 laser engine, beam expanders, reflection mirrors, a three-axis dynamic scanner, and an x-y galvano mirror. The three-axis dynamic scanner system is employed instead of an fθ lens especially for sintering large objects.

The three-axis dynamic scanner system is employed instead of an fθ lens especially for sintering large objects. This scanner system has the ability to prevent spotdistortion by applying a focal distance function when the laser is used to irradiate large areas.This three-axis dynamic scanner contains an objective lens and a concave lens located in front of the x-y galvano mirror. The laser module with dual lasers is thus able to sinter large objects.Fig. 2 Dual laser scanning systemPreparation of nanocomposite powder for the dual laser sintering processSynthesis of polyamide-12 powder As the polymer matrix, polyamide-12 (PA-12) was chosen because of their good thermalstability and mechanical properties. Material properties of polyamide-12 are shown in Table 1.Table 1.

Material properties of polyamide-12 are shown in Table 1.Table 1. Material properties of polyamide-12DensityTensile strength1.01 g/cm3 35 MpaElastic modulusMelting temperatureGlass transition temperature1100 Mpa 179 oC 115 oCTo process a large amount of the PA-12 powder that is smaller 100 μm in diameter, a wetprocess was employed. Compared with dry grinding process, the wet process can introduceadditives to improve both the properties and manufacturability of the powder, where over 95% of particles have diameter less than 100 µm. Furthermore, the wet process can also control theparticle shape.The procedure of the wet process is as follows. PA-12 (50 g) pellets (EMS-Grivory) aredissolved in a mixture of benzyl alcohol (300 g) and Xylene (100 g) in a container where a stirrer and a condenser are installed. A 2% volume fraction of dispersing agent and stabilizer are added to the solution of PA-12.

A 2% volume fraction of dispersing agent and stabilizer are added to the solution of PA-12. The temperature of the solution is increased to 140 ± 5o C for 1 h while being stirred to fully dissolve PA-12. When the solution cools down to 115 ± 2o C, PA-12 begins to crystallize in the solution. When the crystallization ends, the solution is cooled down below 50 oC to separate out the PA-12 crystals by a filter installed in a vacuum pump. The residue solvent in the container is completely evaporated in the vacuum oven. Finally the desired PA-12powder is obtained by sieving the crystallized particles smaller than 100 µm in diameter.1003.5903.0%80/) x ( 3 Q n o i t u b i r t s i d e v i t a l u m u c7060504030202.52.01.51.0) x ( g l 3 q n o i t u b i r t s i d y t i s n e d100.50 468 10204060 80 1002000 400 600 8001000particle size(a) Particle analysis data (b) SEM imageFig.

3 Evaluation of the PA-12 powderFigures 3 (a) and (b) respectively show the particle analysis data and an SEM micrograph ofPA-12 powder manufactured by the above procedure. The average particle diameter is 49.06 μmand the particles have a fairly uniform size. As shown in Fig. 3 (b), particles are almost spherical,the desired shape to ensure uniform layering and precise sintering.Functionalization of graphite nanoplateletsGraphite nanoplatelets (GNPs) are disk-shaped graphite particles of nanometer scalethickness. GNPs are attractive as fillers because they offer high strength, stiffness, and electricaland thermal conductivities. Compared with carbon nanofibers or nanotubes, GNPs providereinforcement in two directions and are expected to be more efficient in improving mechanical properties.As-received GNPs (ASBURY graphite mills Inc., grade 3775) do not have optimuminterfacial bonding in the polymer composites.

However, the interfacial bonding can beimproved by chemically or physically modifying the surface of GNPs. Recent publications show that by functionalizing the surface of GNPs via a nitric acid oxidation process one cansubstantially increase the tensile strength and modulus of the resulting nanocomposite. Thenitric acid oxidation leaves carboxylic acid groups on the surface that chemically bonds with theepoxy [10, 11].The procedure used for the surface functionalization of GNPs is as follows [11]. As-received GNPs were chemically treated with 67% of nitric acid, HNO3, at 83 °C. The platelets were then rinsed with de-ionized water for 2 hours and dried in an oven at 200 °C for 24 hours to removeresidual acid and water. The GNPs produced were approximately 3 μm in diameter and 25 nm in thickness as shown in Fig. 4.Fig.

The GNPs produced were approximately 3 μm in diameter and 25 nm in thickness as shown in Fig. 4.Fig. 4 SEM image of Nitric acid treated GNPsMixing of GNPs and PA-12 powder GNPs were mixed with the PA-12 powder using a Thinky mixer (THINKY CORP., AR200) toachieve uniform dispersion for 1 hour and a sieve shaker (Retsch, AS200) to eliminateagglomerates. The mixer employs cup's dual high speed spinning motions that create forces of over 400 G's to accomplish powerful mixing and de-aerating simultaneously.Figures 5 (a) and (b) show the nanocomposite powder thus produced smaller than 100 µm indiameter. Figure 5 (b) shows GNPs with sharp edges, shown inside the circle, adhering on apolymer particle. The nanocomposite powder used in the dual laser sintering process contains 0.5% weight fraction of GNPs.(a) Nanocomposite particles (b) GNP on a polymer particleFig.

6 Sintering of the nanocomposite in the dual laser sinteringBefore laser sintering begins, the powder was predried at 110 oC for 4 hours to avoid the agglomeration by moisture. In the current test, the powder was sintered in the shape of a tensilespecimen in one layer. Figure 6 shows no sign of agglomeration without irregular melting.Although the result is preliminary, the synthesized nanocomposite powder is thus believed tohave good processability in the SLS process. Work is continuing to demonstrate that the nanocomposite powder can be used to fabricate products with better mechanical and electricalproperties.Conclusions A nanocomposite powder based on graphite nanoplatelets (GNPs) and PA-12 particles has beensynthesized for use in the selective laser sintering process. To improve interfacial bonding, theGNPs were functionalized by a nitric acid oxidation treatment. The two types of particles were mixed in a rotary Thinky mixer.

To improve interfacial bonding, theGNPs were functionalized by a nitric acid oxidation treatment. The two types of particles were mixed in a rotary Thinky mixer. The resulting nanocomposite powder indicates uniform mixingwithout agglomerations. The preliminary result on laser sintering experiment indicates thesuitability of the nanocomposite powder for use in the laser sintering process. Efforts areunderway to confirm better mechanical properties expected of the prototypes made of the nanocomposite powder.AcknowledgementsThis study was sponsored by “Three-Dimensional Digital Real Object Duplication System(RODS)” Development Project from the Ministry of Commerce, Industry and Energy. Weappreciate and thank everyone who has supported and participated in this study.References [1] P. F. Jacobs, Stereolithography and other RP&M Technologies from Rapid Prototyping toRapid Tooling, ASME Press, New York, 1996.[2] D.Y. Yang, H.C. Kim, S.H.

VELOPMENT OF NANOCOMPOSITES FOR SOLID FREEFORM FABRICATIONA. Manthirarn, F. Chi,L. F.Johnson, B. R.Birrningharn andH.L. Marcus Center for Materials Science and Engineering, ETC 9.104 The University of Texas at Austin Austin, TX 78712ABSTRACTNanocomposites in which the constituents are mixed on a nanorneter scale can provide important advantages in the Selective Laser Sintering (SLS) and Selective Laser Reactive Sintering (SLRS) processes. The larger surface area and grain boundaries in the nanocolnposites compared to that in the conventional microcomposites are expected to enhance the solid state diffusion during laser irradiation as well as during any other subsequent processes. Our strategy is to design and develop nanocomposites in which one nanosize cOlnponent has a lower melting point than the other nanosize component, either of which can serve as the matrix phase. The nanoscale dispersion of the low melting component can aid the sintering process during SLS or SLRS.

The nanoscale dispersion of the low melting component can aid the sintering process during SLS or SLRS. Nanocomposite powders of AI203-COOx, Ab03-NiO, A1203-CO and A1203-Ni have been synthesized by sol-gel processing and are evaluated by SLS.INTRODUCTIONNanostructured materials are characterized by ultrasmall building blocks whose sizes are of the order of 1-100 nanometers. These materials offer potentially useful and unusual combinations of mechanical, electrical, magnetic and optical properties that are otherwise unachievable in conventional materials [1]. The properties are controlled by a complex interplay among the building blocks and the interfaces between them that are determined primarily by the synthesis and processing conditions.

The properties are controlled by a complex interplay among the building blocks and the interfaces between them that are determined primarily by the synthesis and processing conditions. Nanostructured materials promise to advance a range of technologies and there is intensive worldwide activity on these materials both from scientific and technological points of view.While the terms "nanocrystalline" and "nanophase" materials refer generally to single phase solids of nanometer size, the term "nanocomposite" refers to multiphase solids of nanometer size [2]. The solid phases in a nanocomposite could be of different nature. They could be metals, inorganic ceramics or polymers or combinations thereof but all mixed on a truly nanometer scale. The phases could be amorphous, semicrystalline or crystalline.Nanocomposites exhibit ultraheterogeneity or nanoheterogeneity, which offer several processing advantages.

The phases could be amorphous, semicrystalline or crystalline.Nanocomposites exhibit ultraheterogeneity or nanoheterogeneity, which offer several processing advantages. Because the different phases are mixed on a nanoscale, the diffusion length for solid-solid reaction is much smaller than that in conventional materials. The larger surface and grain boundary energies associated with nanocomposites provide the driving force for enhanced reactions. These factors reduce the reaction temperature as well as time. Nanocomposites have been shown to sinter close to theoretical densities in several systems such as Ab03-Si02, Ab03-Ti02, Si02-MgO, etc. [3]. Nanocomposite gels store much higher metastable energy than homogeneous gels, and this extra energy enhances64densification. In addition, nanocomposites seem to undergo crystallization and densification silnultaneously unlike homogeneous gels.

In addition, nanocomposites seem to undergo crystallization and densification silnultaneously unlike homogeneous gels. Homogeneous gels, due to their atomic scale mixing, often undergo crystallization before densification occurs which either leads to lower densities or necessitates higher temperatures for sintering. It is clear that nanocomposites have the potential to offer lower processing temperatures and flaw-free parts with higher densities.Our objective is to design and develop nanocomposites that can optimize Selective Laser Sintering (SLS) and selective Laser Reactive sintering (SLRS). Nanocomposites can be synthesized by sol-gel processing or other "soft chemistry" techniques. Although both sol-gel processing and solid freeform fabrication are well known, the idea of utilizing sol gel derived nanocomposite powders that are specifically designed and synthesized to optimize SLS or SLRS have not been exploited.

Our strategy is to select ceramic-ceramic or ceralnic-metal nanocomposite systems in such a way that one nanosize component has a lower melting temperature than the other nanophase component, either of which can serve as the Inatrix phase. The objective is to achieve sintering aid from the low melting phase during SFF processing. Because the low melting component is dispersed on a nanoscale in the matrix phase, it is expected to aid the sintering process, during SFF or any other secondary processing, in a much better way than the mechanically mixed microcomposite powders. By proper selection of composite systems and control of agglomeration during sol-gel processing, the ability to achieve full density parts can be enhanced. We present in this paper our preliminary results with A1203-CoOx, A1203-NiO, A1203-Co and A1203-Ni nanocomposites.

We present in this paper our preliminary results with A1203-CoOx, A1203-NiO, A1203-Co and A1203-Ni nanocomposites. CoO, NiO, Co and Ni have melting points of respectively, 1795, 1984, 1495 and 1455 °C, which are lower than that of Al203 (2050 °C).EXPERIMENTALThe ceramic-ceramic nanocomposite powders of A1203-CoOx, and A1203-NiO were synthesized by a sol-gel method [4]. The steps involved in the synthesis of Al203 coax are summarized in Fig. 1. To an aqueous solution of CO(N03)2. 6H20 maintained at about 80°C, required amount of a solution of aluminum tri-sec-butoxide in iso-propyl alcohol was added under constant stirring. Aluminum-tri-sec-butoxide hydrolyzes during this process to give large suspended particles of Al(OHh, which was then dissolved by adding nitric acid to obtain a homogeneous sol of A13+ and C02+. The sol was heated gently on a hot plate to form a gel, which was then dried rapidly on the hot plate.

The sol was heated gently on a hot plate to form a gel, which was then dried rapidly on the hot plate. The dry gel was then heated in a muffle furnace to decompose the nitrate and hydroxide, which resulted in a mixture of A1203-C0304. A similar procedure was adopted to obtain A1203 NiO also. The A1203-C0304, and A1203-NiO powders were also fired at different temperatures ranging from 300-1250 °c to evaluate the reaction process. The ceramic ceramic nanocomposites so obtained were reduced in a stream of hydrogen at 400-1100 0C to obtain the ceramic-lnetal nanocomposites Al203-Co, and Al203-Ni.The composite powders were characterized by x-ray diffraction, electron microscopy and density Ineasurements. Since the large surface area associated with the nanosize powders lead to an adsorption of water, C02 etc. on the surface, the powders were baked at about 200°C for about 24 h before scanning with the laser beam.

on the surface, the powders were baked at about 200°C for about 24 h before scanning with the laser beam. Single layer tests were performed with various power levels and scan speeds. The specimens obtained were characterized by x-ray diffraction and microscopy.65I Al(OC4H9h in isopropyl alcohol IIC (N03h. 6H20 in water0Mixed under stirring at 80 OCAl(OHh precipitate + Cobalt nitrate solutionPeptization with HN03 at 80 °cAl3+ + C02+ solDry on a hot-pIateI Dry gel I ~ Al20 3 + C0304 powder mixture ICalcination at various temperatureFig.1 Synthesis scheme for Al203-C0304 nanocomposite powderRESULTS AND DISCUSSIONPowder synthesis and characterization:The phase identification by x-ray of a nominal composition containing an atom ratio of Al:Co = 2: 1 is summarized in Table 1. It can be seen that C0304 is formed during the initial stages of drying on the hot plate.

It can be seen that C0304 is formed during the initial stages of drying on the hot plate. Although the exact temperature of hot-plate was not measured, it was presumed that C0304 formation can occur at temperatures as low as 2000C because firing of pure Co(N03h.6H20 at 200°C for 20 h gives C0304. Up to about 4500C, mostly the reflections of C0304 are discernible in the x-ray pattern. Aluminum oxide or AlOOH present are very small in particle size and could not be detected by x-ray diffraction. In the region 580-900 °C, aluminum oxide was found to exist as y-A1203 having a cation-deficient spinel structure. The presence of y-A1203 is, for example, seen in the x-ray pattern (Fig. 2) of a powder mixture fired first at 580°C in air and then at 700 0C in hydrogen atmosphere. A small reflection corresponding to y-A1203 is seen around 28 = 67 ° among strong reflections of Co metal.

A small reflection corresponding to y-A1203 is seen around 28 = 67 ° among strong reflections of Co metal. Above 1000 °C, any aluminum oxide present is in the form of a-Al203 having the corundum structure.66Table I Phase identification by x-ray diffraction of the A1203-Co304 nanocomposites after heating at various temperatures for a constant time of 20 hTemperaturePhase identificationColor(OC)Dried on a hot-plate300 450 580 650 700 800 900 1000 1250C0304 C0304 C0304 C0304 + y-Ah0 3 C0304 + y-Ah0 3 C0304 + y-Ah0 3 C0304 + y-A1203 + CoAh04 C0304 + y-Al203 + CoAl20 4 CoAl204 + a-Al203 CoAl20 4Black Black Black Black Black Black Dark blue Dark blue Dark blue Bluer - - - - - - - - - - - - - - - - - -Co - - - - - ,Co•Co•AI203•78747066625854504642382-Theta in degreeFig.

2 X-ray diffraction pattern of 75 mole % Al203 - 25 mole % C0304 (Al : Co ratio = 2:1) nanocomposite after firing first at 580 OC in air and then at 700 °C in Hz for 20 hThe oxides of cobalt and aluminum can react to give the spinel oxide CoA1204, which is blue in color. The change in color from black to blue around 800 °C (Table 1) seems to suggest that the formation of CoA1204 occurs around this temperature. However, the formation of CoAl204 spinel can start much below 800 °C and we cannot rule out its presence below 800 °C from x-ray data since both C0304 and CoAl204 have spinel structure with close lattice parameters. The broad nature of the reflections at lower temperatures and their overlap make it extremely difficult to determine the exact temperature67at which CoAh04 formation begins. Nevertheless, we could clearly see the formation of the spinel NiAh04 in the analogous system AIZ03-NiO at temperatures as low as 320 °C (Fig. 3).

Nevertheless, we could clearly see the formation of the spinel NiAh04 in the analogous system AIZ03-NiO at temperatures as low as 320 °C (Fig. 3). Formatiqn of spinel phases around 300 °C has also been shownin several other systems such as NiMnz04 [5]. Furthermore, formation ofCoAlzQ4hasbeensuggestedto occur around 500 QC frorn an observation of a decline in the catalytic activity of C0304 supported on "/-AlzQ3 [6]. The catalytic activity in this system is due to. C0304 and the degradation of catalytic activity corresponds to a conversion ofC0304 to CoA1204. So, it is likely that the formation of the CoAlz04 spinel occurs at temperatures below 500 OC.A. • NiO!'I iAI2041000 C.........:: :;;... -<7062544638302· Thetain degreesFig. 3 X-ray diffraction pattern of 50 mole % Al203 - 50 mole % NiO (Al : Ni ratio = 2: 1) nanocomposite after firing at various temperatures for about 20 h6850 nmFig.

3 X-ray diffraction pattern of 50 mole % Al203 - 50 mole % NiO (Al : Ni ratio = 2: 1) nanocomposite after firing at various temperatures for about 20 h6850 nmFig. 4 TEM micrograph of 75 mole % A1203 - 25 mol % C0304 (Co: Al ratio = 1:2) nanocomposite after firing at 580°C in air for 20 hThe TEM micrograph of the A1203-C0304 nanocomposite powder after heating at 580°C in air for 20 h is shown in Fig. 4. The large particles are about 50 nm in size and were presumed to be C0304 since they were shown to be rich in Co by EDS analysis. The small particlesarelessthan 10nm and were shown to be rich in Al by EDS analysis. These observations are consistelltwith the x-ray results. The TEM image of thenanocomposite powder after heating at 700 DC for 20 h is shown in Fig. 5. The particles have grown significantly toabout150nmduring this process. However, the rise in firing temperature did not have any significant influence on the size and distribution of the agglomerates.

However, the rise in firing temperature did not have any significant influence on the size and distribution of the agglomerates. An SEM micrograph of the powder heated at 700 DC is shown in Fig. 6, which indicates an agglomerate size of about 100 Jlm.Selective laser sintering:Preliminary SLS experiments with the A1203-C0304 nanocomposites indicated blowing up ofthe.powder under.the laser beam. A.low powder density of about 20 % together with theadsorbedmoisture<seem to cause this problem. The powder was then ball-milled for about 30 min and subjected to SLSunderanin situ baking around 200 DC. This Inodificationalloweduse of high enough laser power without the powder being blown. Duringthisprocess,the color ofthe powder changedfrom black to blue indicating the reaction ofA1203andC0304underthe laser bealn tofonn the spinel oxide CoA1204. A top view after SLSof.a 83.2 •• mole% .. Al203-16.8mol}%Co304 •• nanocomposite is shown in Fig. 7. •..

A top view after SLSof.a 83.2 •• mole% .. Al203-16.8mol}%Co304 •• nanocomposite is shown in Fig. 7. •.. The corresponding SEM micrograph of the cross section is shown in Fig. 8. Althoughthe>powdersinterslocally during<theSLS process to give relatively higher density, large cracks were found to occur. The low density of the initial powder with large porosity was suspected to cause this problem. Similar results were also obtained with Al203-NiO.6950 nmFig. 5 TEM micrograph of 75 mole % Ah03 - 25 mole % C0304 (Co: Al ratio = 1:2) nanocomposite after firing at 700°C in air for 20 hFig. 6 SEM micrograph of75 mole % Al203 1:2) nanocomposite after firing at 700°C in air for 20 hmole % C0304(CO : Al ratio =707 Photograph of the top view of 83.2 mole % Ab03 - 16.8 mole % C0304nanocomposite after firing at 580°C in air for 20 h and subjecting to SLS.10mTop surface of the layer ~Fig.

8 SEM micrograph of 83.21TIole % nanocomposite after firing at 580°C in air for 20 h and subjecting to SLS.16.8 mole % C030471the density of the initial powders without sacrificing much of the nanoscale mixing were pursued. One way to achieve higher densities in the powder is to annealthe.nanocornpositesatrelatively higher telnperatures (.... 1000 0 C). However, such an annealing around 1000 °c of the A1203-C0304 nanocomposite will result in the formation of the spinel oxide CoAb04, which will in essence consume the low melting component C0304 and jeopardize our approach to SLS. Therefore, to demonstrate the feasibility of our objective, we decided to pursue the ceramic-metal nanocomposite Ab03-CO, which will not undergo any chemical reaction during the preannealing of the powder at higher temperatures. Accordingly, A1203-CO nanocomposites were first obtained by reducing the Ab03-C0304 oxide nanocomposite in hydrogen around 1000 oC.

Accordingly, A1203-CO nanocomposites were first obtained by reducing the Ab03-C0304 oxide nanocomposite in hydrogen around 1000 oC. The powder was then annealed in N2 atm at 1350 °c for 8 h to increase the density, and reduced again in hydrogen around 1000 °c to convert back any cobalt oxide that might have formed by reaction with the residual oxygen during N2 annealing. The resulting Al203-Co powder was found to have a density of about 42 % and sinter under the laser beam. The single layer test coupons appear to have better structural integrity with much reduced macrocrack formation. An SEM photograph of the powder after SLS is shown in Fig.

The single layer test coupons appear to have better structural integrity with much reduced macrocrack formation. An SEM photograph of the powder after SLS is shown in Fig. 9.In order to circumvent the above difficulties, strategies to9 SEM micrograph of 50 mole % Al203 - 50 mole % Co nanocomposite afterfiring at 1350 °C in N2 for 8 h, 1000 0C infor 8 h and subjecting to72CONCLUSIONSNanocolnposite powders of Ah03-COOx, Ah03-NiO, Ah03-CO and Ah03-Nihave been successfully synthesized by sol-gel processing. The powders have been investigated by SLS. Low densities (about 20 %) of the initial powders not only lead to a formation of macrocracks during SLS, but also prevent use of high enough laser power that can cause sintering. Increasing the initial powder densities to about 40 % seems to avoid this difficulty significantly and give structurally sound specimens.

velopment of Nickel-Titanium Graded Composition Components M. S. Domack* and J. M. Baughman** *NASA Langley Research Center, Hampton, VA 23681 **Lockheed Martin Space Operations, Hampton, VA 23681AbstractThe potential of various manufacturing methods was evaluated for producing nickel-titanium graded composition material. The selected test case examined attachment brackets that join nickel-based metallic thermal protection systems to titanium-based launch vehicle structure. The proposed application would replace nickel-based components with graded composition components in an effort to alleviate service induced thermal stresses. Demonstration samples were produced by laser direct metal deposition, flat wire welding, and ultrasonic consolidation.

Demonstration samples were produced by laser direct metal deposition, flat wire welding, and ultrasonic consolidation. Microstructure, general bond quality, and chemistry were evaluated for the components.IntroductionThe ability to fabricate bulk materials with graded composition offers the potential forunique solutions to engineering problems and offers advantages over conventional materials and traditional composites. Research and development of functionally graded materials has largely focused on surface coatings or modifications of surface layers to enhance wear, optical, or electrical properties. Application of functional grading in bulk form offers the ability to create components with tailored chemistries or microstructures which enables properties to be selectively enhanced in critical regions.

The current study examines the potential of various manufacturing methods for producing bulk material with composition graded from nickel-based Inconel 718 to titanium alloy Ti – 6Al – 4V (Ti-6-4). The test case for application of such material is an attachment bracket that joins nickel-based metallic thermal protection systems to titanium-based launch vehicle structure. The goal of the current study was to evaluate the compatibility of each method with the materials selected, examine the resulting microstructures and chemistries, and characterize any reaction products.Metallic thermal protection systems (TPS) under development offer potentially greaterdurability than ceramic tile systems. One design, shown in Figure 1, employs attachment brackets to join a hot service aeroshell to an integrally stiffened tank structure [1, 2]. In this concept the attachment bracket is made from Inconel 718 strip approximately 4 in. long, 1 in. wide, and 0.020 in. thick, with 0.5 in.

In this concept the attachment bracket is made from Inconel 718 strip approximately 4 in. long, 1 in. wide, and 0.020 in. thick, with 0.5 in. long flanges at each end to facilitate brazing to the Inconel 718 aeroshell and the Ti-6-4 tank structure. Thermal analysis indicates that temperatures at the Inconel end would reach 1800oF and be in the range of 400-500oF at the Ti end. The limitation of this design is accommodation of thermal expansion during flight without generating excessive bending or shear loads that impact the service life of the attachment bracket. Replacement of the Inconel 718 attachment clip with one produced from a graded composition component may reduce the stress concentrations.Three manufacturing routes were evaluated for producing a composition graded fromInconel 718 to Ti-6-4 in bulk form: laser direct metal deposition, flat wire welding, and ultrasonic consolidation.

All of these processes are layer-additive but differ in the feedstock and method of energy input used to consolidate material. Each process offers unique manufacturing flexibility and process benefits. Laser direct metal deposition consolidates deposited metal655powder using a high energy laser beam. Flat wire welding is a pulsed laser based method of joining sequential layers of metallic wire. Ultrasonic consolidation is a solid state process that consolidates foil layers by application of frictional energy under load.TPS Attachment ClipFigure 1. Concept drawing of a metallic thermal protection system employing attachment clips to join a hot Inconel 718 aeroshell to a Ti-6-4 integrally stiffened tank structure.Materials and ProcessesLaser direct metal deposition trials were conducted with the Laser Engineered Net Shape(LENS) process developed at Sandia National Laboratories [3].

The LENS process, shown schematically in Figure 2, is a layer additive powder deposition process that consolidates powder layers with a laser operated in continuous beam mode. The LENS process employs multiple nozzles to feed elemental and pre-alloyed powders to produce nearly continuous composition gradients. A wall-shaped build in progress is shown in Figure 3. The deposits for this study were produced on Ti-6-4 base plates from powders of Ti-6-4 and/or Inconel 718. TheLaser and Powder Delivery NozzleX--Y Positioning StageFigure 2. Schematic of the laser engineered netFigure 3. Wall-shaped build in progressshape (LENS) process.using the LENS process.chemistries for the supplied powders are provided in Table 1. Deposits were produced from Ti-6-4 and Inconel 718 in order to establish deposition parameters.

Deposits were produced from Ti-6-4 and Inconel 718 in order to establish deposition parameters. A series of constant composition deposits were produced from pre-mixed powder blends, with the composition adjusted in steps of 10% by volume from 100% Ti-6-4 to 100% Inconel 718, to investigate deposition performance for specific alloy combinations. Deposits with continuous composition656gradients were achieved by simultaneously adjusting the feed rate of each powder in order to achieve a target composition profile and the deposition parameters in order to control the build quality. Deposit geometries included walls, columns, and cylinders.Table 1.

Deposit geometries included walls, columns, and cylinders.Table 1. Chemical compositions of Inconel 718 and Ti-6-4 powder in weight percent.Element Iron Nickel Chromium Columbium Tantalum Molybdenum Titanium Aluminum Vanadium OxygenInconel 718 18.4 52.4 19.1 5.02 <0.004 3.15 1.00 0.32 0.010 0.043Ti-6-4 0.091 0.020 0.0074 <0.002 <0.004 0.0053 88.8 6.29 4.17 0.081Flat wire welding trials were performed using the Precision Metal Deposition (PMDTM) process developed by H&R Technology, Inc. [4]. The PMDTM process, shown schematically in Figure 4, is a pulsed laser-based metal additive fabrication process. The PMDTM process fuses solid flat metal wire onto a substrate without the need to create a molten pool of metal on the substrate. As a result, metal parts can be fabricated with an order of magnitude less heat input than continuous-beam laser metal deposition processes, yielding parts with minimal distortion and improved metallurgical characteristics.

Composition gradients in bulk components can be tailored by adjusting the stacking sequence of wires of different composition. The current study evaluated the compatibility of the PMDTM process with Inconel 718 and Ti-6-4 as well as the interactions that occur when the alloys are combined in a gradient composition component.All of the flat wire welded deposits were produced on Inconel 718 base plates using Ti-6-4 and/or Inconel 718 rectangular cross section wire, as summarized in Table 2, and included three or four layers with three wires deposited side-by-side in each layer. Deposits were produced from Inconel 718 wire, 0.018 in. thick and 0.028 in. wide, in order to establish process parameters. In addition, deposits were produced from Ti-6-4 wire in order to establish process parameters and to investigate interactions with the Inconel 718 base plate. The Ti-6-4 wire used had dimensions 0.016 in. thick and 0.027 in. wide or 0.005 in. thick and 0.028 in. wide.

The Ti-6-4 wire used had dimensions 0.016 in. thick and 0.027 in. wide or 0.005 in. thick and 0.028 in. wide. Graded composition deposits were made by depositing three layers of 0.005 in. thick Ti-6-4 wire followed by one layer of Inconel 718 wire.Ultrasonic consolidation (UC), developed by Solidica, Inc., is a micro-friction solid-stateprocess used to join foil gage feedstock to build parts in a layer additive fashion [5]. The UC process, shown schematically in Figure 5, produces metallurgical bonds between layers without melting. The layer being added is translated against the previous build at ultrasonic frequency and low amplitude resulting in fracture and displacement of surface oxides. The clean surfaces657are brought together at a temperature typically less than half the material melting temperature and under applied load to promote atomic diffusion. The sonotrode is rotated across the build surface to continuously consolidate foil strip.Table 2.

The sonotrode is rotated across the build surface to continuously consolidate foil strip.Table 2. Materials and Configuration used forInconel 718 and Ti-6-4 Deposits.Sample NumberAlloy DepositedSubstrate MaterialNumber of LayersWires per LayerWire Dimensions Thickness (in.)Width (in.)1-1In 718In 718330.0180.0281-2In 718In 718330.0180.0281-3In 718In 718330.0180.0282-1Ti-6-4In 718330.0160.0272-2Ti-6-4In 718330.0050.028Figure 4. Schematic of the Precision Metal Deposition (PMDTM) Process.2-33-1Ti-6-4Ti-6-4 + In 718 (Layer4)In 718In 71834330.0050.005 + 0.0180.0280.0283-2Ti-6-4 + In 718 (Layer4)In 718430.005 + 0.0180.028TravelRotating ultrasonic consolidation headLayer being addedSonotrodeSubstrate (previously deposited layers)Instantaneous deformation affected zoneFigure 5.

Schematic of the ultrasonic consolidation process.Nickel and titanium alloys are suitable for ultrasonic consolidation, but the process is notmature for these materials. Samples were produced for this study from pure Ni and commercially pure (CP) Ti foils to demonstrate the feasibility of developing metallurgical bonding without undesirable metallurgical interaction. For this feasibility study, static joining of small 0.24 in. x 0.6 in. samples was attempted by using a modified spot welding system equipped with a knurled tool steel sonotrode, rather than the rolling strip application. Samples were658produced from pure Ni and from CP Ti in order to establish processing parameters. Graded samples were produced by layering Ni and CP Ti foils in the sequence Ni-Ni-Ni-Ti-Ni-Ti-Ti-Ti. The CP Ti foil was 0.002 in. thick rolled foil, and the Ni foil was 0.002 in.

Graded samples were produced by layering Ni and CP Ti foils in the sequence Ni-Ni-Ni-Ti-Ni-Ti-Ti-Ti. The CP Ti foil was 0.002 in. thick rolled foil, and the Ni foil was 0.002 in. thick electrodeposited foil.Analysis of all samples produced for this study employed standard metallographictechniques for sectioning and polishing. However, etching procedures were tailored for each product and material combination in order to examine bond lines, evaluate material interactions, and reveal grain structures. Chemical analysis was evaluated by direct current plasma spectroscopy for bulk samples and by energy dispersive x-ray analysis of specimen sections.Results and DiscussionLaser Direct Metal DepositionTest samples produced by the LENS process for this study are shown in Figure 6. Duringbuilding of the Ti-6-4 wall (2.5 in. x 4.0 in.) power was adjusted to control rippling and smooth the surface.

Duringbuilding of the Ti-6-4 wall (2.5 in. x 4.0 in.) power was adjusted to control rippling and smooth the surface. Power and feed rate were adjusted during building of the series of Inconel 718 walls (2.5 in. wide x 2.5 in. tall max.) in order to reduce rippling, warping, and separation from the base plate. The Inconel 718 deposits generally had a smoother surface than that of the Ti-6-4 deposits. Metallographic sections of the Inconel 718 wall, shown in Figure 7 etched with a solution of HNO3 and HCl, exhibited banding associated with the laser passes. The microstructure was primarily a fine scale dendritic solidification structure with coarser dendrites noted near the surface of the deposit. The Ti-6-4 wall, shown in Figure 8 etched with Keller’s reagent, exhibited significant surface roughness characterized by serrations that were on the scale of the deposited powder layer. Residual powder particles were loosely adherent to the surface of the deposit.

Residual powder particles were loosely adherent to the surface of the deposit. The microstructure exhibited large columnar grains, prior-β grain boundaries, and fine alpha laths within the grains. The columnar grain structure indicates that epitaxial growth of β likely occurred during laser consolidation of the added powder layer with transformation of β to α during cooling.Ti-6-4InconelInconelComposition BlendsGradientGradientGradientFigure 6. LENS samples produced from Ti-6-4 and Inconel 718 powders.659(a)(b)(c)500 µm100 µm100 µmFigure 7. Microstructure of the Inconel 718 deposit illustrating (a) laser pass boundaries,(b) fine dendrite structure, and (c) coarse dendrite structure.(a)(b)1000 µm200 µmFigure 8. Microstructure of the Ti-6-4 deposit illustrating (a) surface roughness, large columnargrain structure, and (b) alpha laths.Constant composition blends of the two materials were deposited to build walls 1 in.wide x 0.5 in.

tall and columns 0.375 x 0.375 x 0.5 in. tall. Cracking was observed in the walls when the composition blend was 50% each Ti-6-4 and Inconel 718 and in the columns when the blend was 40% Ti-6-4 and 60% Inconel 718. Chemical analysis of the constant composition blend column samples is summarized in Table 3 for the various powder blends. The results indicated that the deposits targeted to be 100% Ti-6-4 and 100% Inconel 718Table 3.

The compositions of the remaining powder blends differed by 10-20% from calculated compositions based on target powder volume percent, suggesting that further development is needed in powder feed control. The deposits with target powder blend of 60-80% Inconel 718 agreed best with calculated composition. X-ray chemical analysis of the short column deposit with target blend of 10% Ti-6-4 and 90% Inconel 718, summarized in Figure 9, revealed a complex solidification structure with significant elemental segregation. The backscattered scanning electron microscope image shows a coarse dendrite structure with partitioning of nickel and iron, circular aluminum and chromium rich inclusions, and occasional titanium rich inclusions.10 µmTiAlNiCrFeFigure 9.

X-ray chemical maps of the column deposit with target powder blend of 10% Ti-6-4and 90% Inconel 718.The graded composition deposits developed macroscopic cracks (Figure 10) before thefull transition from Ti-6-4 to Inconel 718 was achieved. The 2.5 in. wide wall developed a large crack when the composition reached a target mix of approximately 58% Ti-6-4 and 42% Inconel 718, based on powder feed control, with the crack propagating to complete fracture of the build during shutdown and removal of the sample. The column (0.375 in. x 0.375 in. x 0.98 in.) developed a crack when the target blend was about 40% Ti-6-4 and 60% Inconel 718. The build was continued to completion but additional cracks developed. Both deposits exhibited(a)Ti base plate100% Ti-6-4(b)100% IN 718Metallurgical sectionCrack58% Ti-6-4 42% IN 718100% Ti-6-4Ti base plateFigure 10.

Both deposits exhibited(a)Ti base plate100% Ti-6-4(b)100% IN 718Metallurgical sectionCrack58% Ti-6-4 42% IN 718100% Ti-6-4Ti base plateFigure 10. Graded composition (a) wall and (b) column deposits illustrating macroscopic cracks.661significant secondary cracking as well. The microstructures were similar for the column and wall and revealed that the cracks were not directly linked to metallurgical features. Macroscopic banding was observed, as shown in Figure 11a, likely due to the laser passes used to consolidate each powder layer. The sample was etched with an HNO3/HCl solution + Keller’s reagent. The microstructures were fully dendritic with coarser structure in the lighter bands. Residual Ti-6-4 powder particles that were not melted during processing were observed near the surfaces. The primary fracture surface of the wall (Figure 12) exhibited brittle, cleavage fracture features and residual Ti-6-4 powder particles.

The primary fracture surface of the wall (Figure 12) exhibited brittle, cleavage fracture features and residual Ti-6-4 powder particles. A reaction layer was visible around the Ti-6-4 powder particles both in cross section (Figure 11c) and on the fracture surface (Figure 12c).(a)(b)(c)1000 µm100 µm25 µmFigure 11. Metallurgical section of the graded composition wall deposit illustrating (a)microstructure banding, (b) dendrite structure, and (c) residual Ti-6-4 powder.(a)(b)(c)100 µm10 µm25 µmFigure 12. Fracture surface of (a) the graded composition wall deposit illustrating (b) brittlefracture features and (c) exposed residual Ti-6-4 powder.Flat Wire WeldingSamples produced by the PMDTM flat wire welding process, shown in Figure 13,exhibited distinct wire layers (Figure 13a) and solidification fronts (Figure 13b) which were associated with the laser passes along the three wires in each layer.

The Inconel 718 deposits were well integrated with the base plate but the Ti-6-4 and Ti-6-4/Inconel 718 deposits exhibited macroscopic cracking and separation at the base plate interface (Figure 13c). The Ti-6-4/Inconel 718 deposits also exhibited cracking through the layers of the deposit.Metallurgical sections of the Inconel 718 deposit (Figure 14 etched with Kalling’s reagent) showed full consolidation between the wire layers and banding associated with the laser passes. The microstructure exhibited large elongated grains that spanned the laser pass boundaries and were generally elongated in the through-thickness direction (Figure 14). Occasional porosity was also observed. A region of mixing was observed at the base plate interface that penetrated up to 300 µm into the base plate.662(a)Layers(b)Wires(c)Crack50µmBase plate200µmBase plate50µmBase plateFigure 13.

Typical appearance of the flat wire welded samples, illustrating (a) three layers, (b)solidification fronts along the three wires in each layer, and (c) cracking at the base of the Ti-6-4 deposits.(a)(b)Base plate200µmBase plate200µmFigure 14. Metallurgical (a) cross and (b) longitudinal sections of Inconel 718 deposit built onInconel 718 base plate showing solidification fronts, grain structure, and porosity.Metallurgical sections of the Ti-6-4 deposit (Figure 15 etched with Kalling’s + Flick’sreagents) exhibited distinct boundaries between wire layers with variations in the microstructure of each layer. A nearly continuous crack was observed at the base plate interface with numerous secondary cracks through the first Ti-6-4 layer. The bottom Ti-6-4 layer was reduced in thickness compared with the second and third layers and also exhibited regions of mixing with the Inconel 718.

The bottom Ti-6-4 layer was reduced in thickness compared with the second and third layers and also exhibited regions of mixing with the Inconel 718. A reaction layer was observed at the base plate interface, appearing as a 5 µm thick band, bordered by a 25-100 µm thick brittle fracture region.(a)(b)Brittle fracture regionLayersReaction layerCrackBase plate200 µm200 µmBase plate25µmFigure 15. Metallurgical section of Ti-6-4 deposit built on Inconel 718 base plate showing (a)distinct wire layers, cracking, (b) the reaction layer and brittle fracture region.The Ti-6-4 / Inconel 718 deposits exhibited more significant cracking at the base platewith some separation of the deposit, as shown in Figure 16. Metallurgical sections revealed distinct boundaries between the top Ti-6-4 wire layer and the Inconel 718 wire layer but the663boundaries between individual Ti-6-4 layers were not as clear.

Metallurgical sections revealed distinct boundaries between the top Ti-6-4 wire layer and the Inconel 718 wire layer but the663boundaries between individual Ti-6-4 layers were not as clear. Some evidence of mixing as well as secondary cracking was observed in the lower portion of the Ti-6-4 region. Additional mixing was observed at the interface between the Ti-6-4 layers and the Inconel 718 layer. Reaction layers that were approximately 5 µm thick (Figure 17) were observed at the interface between the Ti-6-4 layers and both the Inconel 718 base plate and wire. The reaction layer was similar in appearance to that observed in the Ti-6-4 deposit.IN 718 layer200 µmTi-6-4 layersCrackBase plateFigure 16.

The reaction layer was similar in appearance to that observed in the Ti-6-4 deposit.IN 718 layer200 µmTi-6-4 layersCrackBase plateFigure 16. Metallurgical section of Ti-6-4 / Inconel 718 deposit built on Inconel 718 base plate.Section was etched with Kalling’s and Kroll’s reagents.(a)IN 718 layer(b)IN 718(c)Ti-6-4Ti-6-4 layersCrackReaction layerReaction layer25 µmTi-6-425 µmBase plate100 µmBase plateFigure 17. Metallurgical section of Ti-6-4/Inconel 718 deposit illustrating (a) cracking at thebase plate interface and reaction layers between the Ti-6-4 wire and (b) the Inconel 718 wire and (c) the base plate. Section was etched with Kalling’s reagent.At higher magnification the reaction layer appeared to be a nearly featureless band, approximately 5 µm thick, as shown in Figure 18. The reaction layer shown has formed around a lobe of Inconel 718 drawn into the Ti-6-4 layer during interlayer mixing.

The reaction layer shown has formed around a lobe of Inconel 718 drawn into the Ti-6-4 layer during interlayer mixing. Chemical analysis by energy dispersive spectroscopy was performed stepwise along a line spanning from Inconel 718 at the interior of the lobe to the surrounding Ti-6-4 layer. Semi-quantitative analysis demonstrated a gradient decreasing in Ni, Fe, and Cr and increasing in Ti and Al, but was not sensitive enough to fully quantify the chemistry. Crystallographic analysis and fully quantitative x-ray analysis of the reaction layer are underway and not included in this report.The flat wire welding process was used to produce metallurgical mixing and bondingbetween Ti-6-4 and Inconel 718, although some cracking and material reaction occurred, which indicates the need for further process development.

The pulse application of the laser in flat wire welding resulted in minimal segregation of elements as compared with the LENS process.664InconelTiInconel1 (cid:120)(cid:120)Reaction LayerChemical Analysis(cid:120) (cid:120) (cid:120) (cid:120) (cid:120) (cid:120) (cid:120) (cid:120)10Ti)% wt(n o i t a r t n e c n o C504030201001Ni Al Nb Ti Cr Fe23456789Point on lineFigure 18. Detailed morphology and chemical analysis of a reaction layer formed in the regionof mixing at the Ti-6-4 and Inconel 718 wire interface.Ultrasonic ConsolidationThe samples produced by ultrasonic consolidation exhibited indentations from theknurled surface of the sonotrode in the consolidated region, as shown in Figure 19. The free edges of the foil were flared due to peel loads manually applied by the manufacturer to ensure that the foils were bonded.

The free edges of the foil were flared due to peel loads manually applied by the manufacturer to ensure that the foils were bonded. The samples were sectioned along the length and across the remaining width as shown in Figure 19. The longitudinal section (Figure 20) revealed that there were residual indentations from the knurled surface of the sonotrode at each foil layer which prevented complete consolidation of the materials. The residual indentations are an experimental artifact associated with the spot weld application of the sonotrode. In the commercial process these voids are normally filled by plastic flow of material resulting from the rotation of the sonotrode. Regions of metallurgical bonding were identified on each sample for analysis.Consolidated RegionCross SectionLongitudinal Section0.1 in.Figure 19.

Regions of metallurgical bonding were identified on each sample for analysis.Consolidated RegionCross SectionLongitudinal Section0.1 in.Figure 19. Typical ultrasonically-consolidated sample illustrating knurled surface in theconsolidated region and flared foil edges.Metallurgical bonding was achieved in samples produced from Ni and CP Ti foils. Inboth cases the microstructure of the foils remained unchanged by the consolidation process. The scanning electron microscope cross section image shown in Figure 21a includes five layers within a sample of ten Ni foil layers. The sample shown is in the as-polished condition. The triangular voids are residual indentations from the sonotrode. Metallurgical bonding (MB) is noted at the angled arrows and shown in greater detail in Figure 21b, which has been etched with66510Kroll’s reagent. The metallurgical bond line remained tightly closed after etching and exhibits some smearing of metal across the boundary.

The metallurgical bond line remained tightly closed after etching and exhibits some smearing of metal across the boundary. The Kroll’s reagent also highlighted a centerline feature (C in Figure 21b) which is a remnant of the electrodeposition process used to produce the Ni foil.1000 µmFigure 20. Longitudinal section illustrating residual voids from the knurled surface of thesonotrode. The example shown contains ten Ni foils in the as-polished condition.(a)(b)MBMBMBAMBCFigure 21. Metallurgical cross section of (a) an ultrasonically-consolidated sample containingten Ni foils and (b) morphology of the bond line.A portion of a UC sample containing eleven layers of CP Ti is shown in Figure 22 and has been lightly etched with a solution of HF, HNO3, and lactic acid. Metallurgical bonds are apparent between three of the layers with no bonding at the triangular void.

Metallurgical bonds are apparent between three of the layers with no bonding at the triangular void. Higher magnification (Figure 22b) reveals that the etchant lightly attacked the bond line but the bond remained intact. In addition this view illustrates that the equiaxed alpha microstructure of the foil was unchanged during the consolidation process.(a)(b)Figure 22. Metallurgical cross section of (a) an ultrasonically-consolidated sample containingeleven CP Ti foils and (b) morphology of the bond line.666The graded composition sample, shown in Figure 23, contained both Ni and Ti foils inthe sequence Ni – Ni – Ni – Ti – Ni – Ti – Ti – Ti with the sonotrode applied between each layer. In the scanning electron microscope image shown in Figure 23a the Ni foils appear lighter than the Ti foils.

In the scanning electron microscope image shown in Figure 23a the Ni foils appear lighter than the Ti foils. Metallurgical bonding occurred between some layers but also evident are residual indentations from the sonotrode surface, occasional foil tearing, and separation of the topmost Ti foil. Higher magnification examination (Figure 23b) revealed that metallurgical bonding did occur between Ni and Ti foils, and was also observed between Ni-Ni and Ti-Ti pairs. In addition, the microstructure of each alloy appeared unchanged by the UC process.(a)(b)NiTiTi Ti Ni Ti Ni Ni NiTiNiNi-Ti BondFigure 23. Metallurgical cross section of (a) an ultrasonically-consolidated sample containing Niand CP Ti foils and (b) morphology of the bond line.Higher magnification examination of the bond line between Ni and Ti foils revealed areaction layer that was approximately 1 µm thick, as shown in Figure 24.

Chemical analysis of the reaction layer and the surrounding Ni and Ti foils was performed by energy dispersive x-ray analysis in a scanning electron microscope and indicated (points 2-5) the presence of both Ni and Ti as well as higher concentrations of C and O than in the foils. Additional chemical analysis and x-ray crystallography are underway to determine the composition and structure of the reaction layer.The ultrasonic consolidation process successfully created a metallurgical bond betweenCP Ti and pure Ni without melting either material. The extent and nature of material interaction requires further characterization but is clearly reduced compared with the flat wire welding process.AnalysisCNOTiNi6Ti1 2 3 4 5 61.9 1.3 2.6 4.9 5.7 14.1 21.4 25.4 4.9 13.6 30.1 19.4 5.6 17.7 30.6 23.4 6.3 7.2 65.9 1.5 3.1 10.4 19.7 64.689.3 33.5 32.0 22.7 19.2 2.114253NiFigure 24.

Chemical composition of the reaction layer in weight percent.667Concluding RemarksThe potential of three manufacturing methods was evaluated for producing bulk materialwith composition gradient from Ti-6-4 to Inconel 718. The methods were laser metal powder deposition, flat wire welding, and ultrasonic foil consolidation. All of the manufacturing methods require further development and additional characterization of produced samples before gradient composition materials can be reliably produced.Graded composition samples produced by laser direct metal deposition exhibited significant macroscopic cracking for powder blends that contained 40-60% Inconel 718. Microstructures exhibited coarse dendrites and significant elemental segregation. The chemistries of deposits of fixed powder blends varied in accuracy by up to 20% compared with calculated target chemistry.

The chemistries of deposits of fixed powder blends varied in accuracy by up to 20% compared with calculated target chemistry. Additional development of process parameters and powder feed control are necessary to ensure that target chemistry gradients are achieved without excessive material reactions.The flat wire welding process resulted in layered deposits of Ti-6-4 and Inconel 718 wirethat exhibited good mixing between the wire layers but also exhibited cracking at the alloy interface. A 5 µm thick reaction layer and region of brittle cracking were observed, but the degree of material reaction was significantly reduced compared with laser direct metal deposition.Samples produced by ultrasonic consolidation demonstrated metallurgical bondingbetween pure Ni and commercially pure (CP) Ti foils without melting. Analysis of the bonded regions revealed a 1 µm reaction layer that was thinner than those observed for the other processes evaluated.References1. Blosser, M.

velopment of Novel EBSM System for High-Tech MaterialAdditive Manufacturing ResearchC. Guo1, 2, 3, F. Lin1, 2, 3, W. J. Ge1, 2, 3 and J. Zhang1, 2, 31. Department of Mechanical Engineering, Tsinghua University, Beijing 100084; 2. Key Laboratory for Advanced Materials Processing Technology, Ministry of Education of China, Tsinghua University, Beijing 100084; 3. Biomanufacturing and Rapid Forming Technology Key Laboratory of Beijing, Tsinghua University, Beijing 100084. AbstractElectron beam is more appropriate for metal additive manufacturing (AM) than laser because of its high energy converting efficiency and high absorption for various materials. It becomes a preference for AM study of high-tech material with high melting point, high brittleness or graded material. A novel electron beam selective melting (EBSM) system with dual-material processing capability has been developed in Tsinghua University to meet the wide high-tech material AM research requirement.

A novel electron beam selective melting (EBSM) system with dual-material processing capability has been developed in Tsinghua University to meet the wide high-tech material AM research requirement. A vibration driven powder supplier was developed and the supplier had a high compatibility to various powders. A stable supplying rate and a supplying accuracy less than 7.5% were obtained with the supplier. Two powders can be supplied individually to obtain a mixture with tailored proportion for each powder layer. The mixture is homogenous and the actual proportion is close to the desired value. In order to prolong the spreading comb’s lifetime and avoid tooth breaking, a low deformation powder spreading device was designed based on dual inclined combs and a one-way scraping mechanism. The system provides exchangeable building tanks with sizes of 100 × 100 × 100 mm3 and 250 × 250 × 250 mm3, which can save powder when the part is small and the powder is expensive.

The system provides exchangeable building tanks with sizes of 100 × 100 × 100 mm3 and 250 × 250 × 250 mm3, which can save powder when the part is small and the powder is expensive. The novel EBSM system is capable of building parts with single material and has a potential of dual-material processing.1. IntroductionElectron Beam Selective Melting (EBSM) is an additive manufacturing technique that builds metal parts layer by layer by using an electron beam to selectively scan and melt the powder layer in a vacuum chamber. Compared to the laser process such as Selective Laser Sintering/Melting (SLS/SLM), EBSM has some features. First, the converting efficiency from electric energy to electron beam energy and the absorptivity of electron beam for metal powder are high. As a result, the electron beam can produce highly dense parts with high melting point materials, e.g. Ti-6Al-4V.

As a result, the electron beam can produce highly dense parts with high melting point materials, e.g. Ti-6Al-4V. Second, before melting the cross-section of the part, the electron beam preheats the powder bed to a high temperature with a fast scanning velocity. The preheat step not only slightly sinters the powder so as to avoid the collapse of the lowly conductive powder bed,but also plays a role as the heat treatment, thus efficiently reduces the thermal stress during fabrication and decreases the brittleness of the fabricated parts. Therefore, EBSM is appropriate for manufacturing parts with high brittleness materials, e.g. TiAl alloy. With the above advantages, EBSM was attracting more and more attentions these years and is regarded as a preferential additive manufacturing technique for high-tech metal.Currently, the corporation ARCAM AB has developed several commercial EBSM machines which are increasingly used in medical implant and aerospace industries.

Parts with various materials has been successfully manufactured via EBSM, e. g. 316L stainless steel, Ti alloy, Co-based alloy, Ni-based superalloy, TiAl alloy [1-5]. However, the application of EBSM still needs to be further expanded so that the overall potential of EBSM can be used to its full capacity. One of the possible new applications of EBSM is the manufacturing of graded materials. Some researchers investigated the feasibility of multi-material processing of SLM. Liu, et al. fabricated 316L stainless steel and UNS C18400 Cu alloy dual-material samples with SLM: after depositing many layers of 316L stainless steel, Cu alloy was sequentially deposited on the previous layers. The interfacial characterization of the samples was investigated and it is found that cracks occasionally appear on the interface [6]. Terrazas, et al. fabricated multi-material structures using Ti-6Al-4V and copper powder with the EBSM technology.

Terrazas, et al. fabricated multi-material structures using Ti-6Al-4V and copper powder with the EBSM technology. The fabricated part with the first material was machined to provide a surface on which the second material was deposited [7]. However, the process can not be completed in a single machine run and the interfacial region was narrow, which would affect the performance of the graded material.In addition, the materials for EBSM need further expansion and it calls for a wide material compatibility of the EBSM system. Especially in the material study, the EBSM system is expected to fabricate samples with various materials with various particle shapes and various particle sizes. EBSM system with wide material compatibility would help the researchers to study new material, optimize the process, and get deep insight into the EBSM technology.In this paper, a novel EBSM system with a capacity of dual-materials processing and with a wide material compatibility was developed.

A vibration driven powder supplier was designed. Two different powders can be supplied individually and precisely and then mixed. The mixing ratio of each powder layer can be tailored so that parts with material in gradual change can be fabricated. In order to prolong the spreading comb’s lifetime and avoid breaking of the comb tooth, a low deformation powder spreading device with dual inclined combs and a one-way scraping mechanism were designed. The new EBSM system provides two exchangeable building tanks with sizes of 100 × 100 × 100 mm3 and 200 × 200 × 200 mm3, which can save powder when the part is small and the powder is expensive. Some experiments and their results are presented in this paper.2. Description of the System2.1. Vibration Driven Powder Supplier and the Powder MixerFig. 1 is a schematic of the vibration driven powder supplier.

Description of the System2.1. Vibration Driven Powder Supplier and the Powder MixerFig. 1 is a schematic of the vibration driven powder supplier. A vibration plate is placed at the lower opening of the powder storage, and the space between the vibration plate and the baffle board provides the powder an exit. While the vibrator stops vibrating, the powder forms a repose angle due to internal friction. While the vibrator starts working, the vibration plate moves backand forth at a high frequency. As a result, the powder continuously moves on the vibration plate, generating a stable powder flow off the vibration plate. The powder supplier driven by vibration has no rotatable parts, thus has high reliability in the dust environment.Fig. 2 is a schematic of the dual powders mixer. Two powder suppliers are placed side by side, facing each other.

2 is a schematic of the dual powders mixer. Two powder suppliers are placed side by side, facing each other. A powder mixing box is placed right below the edges of the vibration plates, and a weight senor is used to measure the weight of powder in the mixing box in real time. Once the powder weight reaches the desired value, the vibrator stops working. The two powders can be supplied into the mixing box individually and precisely to obtain a mixture with desired proportion. Then, the mixing box rotates back and forth, blends the two powders, and finally pours the mixture onto the working platform.(a) Stop vibrating(b) VibratingFig.1. Schematic of the vibration driven powder supplier 1, powder storage; 2, baffle board; 3, powder exit; 4, vibration plate; 5, mixing box; 7, working platform; 8, supporting block; 9, weight sensor; 10, connection; 11, Vibrator.Powder APowder B(a) Dual powder supplier(b) Powder mixingFig.2.

Schematic of the dual powders mixer 1, powder storage; 2, baffle board; 3, powder exit; 4, vibration plate; 5, mixing box; 7, working platform; 8, supporting block; 9, weight sensor; 10, connection; 11, Vibrator.2.2. Low Deformation Powder Spreading DeviceDuring the EBSM, balling phenomena often occur, resulting in bumps on the deposited surface. On this occasion, the powder spreader needs to scrap over the bump, leading to the bending of the comb teeth. Vertical combs are commonly used in the additive manufacturing technologies based on powder bed. Unfortunately, permanent deformation or breaking of the comb teeth often occurs. The picture in Fig. 3 shows an example of the permanent deformation of the comb teeth. After scraping over the balling surface several times, irrecoverable deformation is found in some of the comb teeth. The deformed comb leads to an uneven surface of the powder bed, and worsens the fabrication quality in the following deposition.

The deformed comb leads to an uneven surface of the powder bed, and worsens the fabrication quality in the following deposition. In this study, instead of the vertical comb, the inclined comb is utilized so as to reduce the risk of teeth deformation or breaking.Comb toothBumpVertical combInclined combFig. 3. Vertical and inclined combIn order to enhance the reliability of spreading device, a dual-combs design is adopted in the EBSM system. Two parallel combs are placed in tandem, the front comb keeps off the working platform a small distance, while the back one is exactly on the working platform. In addition, one comb is offset half of teeth width relative to the other comb. As a result, powder leaked between the teeth of the front comb can be leveled down by the back comb.Scraping directionView from A1/2 teeth widthScraping directionFig. 4. Dual inclined combsAs shown in Fig. 4, since the comb is inclined, the comb should move towards the inclined direction.

4. Dual inclined combsAs shown in Fig. 4, since the comb is inclined, the comb should move towards the inclined direction. If the comb moves in the opposite direction, the comb teeth have high risk of getting stuck on the uneven deposited surface. Therefore, a one-way scraping mechanism as shown in Fig. 5 is designed. As illustrated in Fig. 5a, by twisting the adjusting screw, the wheel is constrained on the lower surface of the guide rail. As the slide moves forwards, the combs scraps on the working platform and spreads the powder into the building tank. As the slide moves to the end, the wheel rises along the wedge block, lifts the bridge plate, and then moves onto the uppersurface of the guide rail (see in Fig. 5b). Consequently, as the slide moves backwards, the combs returns back to the original position, without contacting with the working platform (see in Fig. 5c).(a) Original position(b) Moving forwards(c) Moving backwards Fig. 5.

5c).(a) Original position(b) Moving forwards(c) Moving backwards Fig. 5. Schematic of the one-way scraping mechanism 1, slider; 2, guide rail; 3, bridge plate; 4, wedge block; 5, working platform; 6, adjusting screw; 7, mounting plate; 8, combs; 9, wheel; 10, building tank.2.3. Exchangeable Building Tanks with Different SizesThe EBSM system provides two exchangeable building tanks with sizes of 100 × 100 × 100 mm3 and 250 × 250 × 250 mm3. As illustrated in Fig. 6, if the part needs to be fabricated is small and the powder is expensive, the larger building tank can be replaced by the smaller one to reduce the material cost.Smaller building tankLarger building tankFig. 6. Exchangeable building tanks3. Experiments and Results3.1. Material Compatibility of the SupplierVarious powders including sponge Ti, gas and water atomized 316L stainless steel and gas atomized Ti-6Al-4V were tested on the powder supplier.

Material Compatibility of the SupplierVarious powders including sponge Ti, gas and water atomized 316L stainless steel and gas atomized Ti-6Al-4V were tested on the powder supplier. For these 6 kinds of powders, the powder supplying was continuous and stable, and their supplying rates were listed in Table 1. The powder supplier has a high compatibility to different powders. The supplying rate increases with the increase of powder flowability. The 300 mesh sponge Ti powder has the lowest flowability, thus its supplying rate is the lowest. The 100 mesh gas atomized 316L stainless steel powder has the highest flowability, and thus has a fastest supplying rate. Among the gas atomized 316L stainless steel powders, the supplying rate increases with the decrease of mesh number. In other words, a larger powder particle size leads to a higher flowability and supplying rate. Another issue influencing the supplying rate is the powder exit size.

In other words, a larger powder particle size leads to a higher flowability and supplying rate. Another issue influencing the supplying rate is the powder exit size. A larger exit size leads to a higher supplying rate. For powder with higher flowability, the powder exit size should be reduced to obtain the same supplying rate.The errors between the desired and actual weight in powder supplying were measured using the 200 mesh gas atomized 316L stainless steel powder, and the results were presented in Fig. 7. The actual weight of powder fluctuates near the desired value, and the maximum error is 7.5%.Table 1.

7. Errors between the desired and actual weight in powder supplying3.2. Homogeneity of Powder MixturePure Cu and Ni powders were used to test the powder mixing quality. The two kinds of powders were mixed by the system at a desired proportion, and then the mixture was spread into a layer. Since the color of Cu powder (red) is quite different from that of Ni powder (gray), an image detection method was developed to measure the actual proportion. Micrographs were taken at 10 different locations on the powder layer. The number of red pixels in the micrograph was computed and the proportion of Cu equaled the ratio of the red pixels number to total pixels number.From the results listed in Table 2, the variation of proportion at different sampling pointsis minor and the actual proportion is close to the desired value, showing a good quality of the powder mixing of the system.Table 2.

Proportion of Cu at different sampling pointsSampling point numberDesired value: 0.1Desired value: 0.3Desired value: 0.5Desired value: 0.7Desired value: 0.910.09740.29810.50910.69260.923320.10040.28730.49910.71890.910530.09460.30890.55540.73450.898740.09530.28540.53640.70760.895250.09390.30700.50910.69750.896160.09490.28430.51450.70120.895770.09470.28650.56030.70060.904380.09410.30370.51280.68030.902290.09170.28680.50210.70140.9126100.09840.29680.52980.70330.9039Average0.09550.29450.52290.70380.9043Variance0.00240.00910.02060.01390.00853.3. Stress on the Spreading Comb ToothA finite elements model was established in ANSYS to simulate the stress distribution on a comb tooth which is scarping over a bump. Fig.8 shows the Mises stress distribution on a bended tooth, and the maximum stress appears at the root of the tooth.

Fig.8 shows the Mises stress distribution on a bended tooth, and the maximum stress appears at the root of the tooth. The maximum Mises stress vs. bump height curves of a comb tooth at various angles were computed and plotted in Fig. 9. The results show that the maximum Mises stress increases with the increase of bump height and decreases with the increase of inclined angle. Compared to a vertical comb, an inclined comb suffers less stress while scarping over a bump with same height. At a bump height of 0.6 mm, the maximum Mises stress of the comb with an inclined angle of 15° is 31.8% of that of vertical comb. A lower stress level means lower risk of breaking for the comb tooth, and thus would efficiently extend the comb lifetime.Maximum mises stressFig. 8.

A lower stress level means lower risk of breaking for the comb tooth, and thus would efficiently extend the comb lifetime.Maximum mises stressFig. 8. Mises stress distribution on a bended toothInclined angle 10°Mises应力 (MPa)h1 (mm)0.00.20.40.60.81.01.21.41.61.82.02.20200400600800100012001400Inclined angle 15°Inclined angle 5°Inclined angle 20°Inclined angle 0°Inclined angle 25°) a P MInclined angle 30°(Inclined angle 35°s s e r t S s e s iInclined angle 40°M m u m i x a MBump height (mm)Fig. 9. Maximum Mises stress of the comb tooth scarping over a bump3.4. Single-Material ProcessingSingle-material processing was conducted on the EBSM-250 system which adopts the above powder supplier and spreading devices. Materials including 316L stainless steel, Ti-6Al-4V were successfully fabricated. The average relative density of fabricated samples reaches up to 99.96%.

Materials including 316L stainless steel, Ti-6Al-4V were successfully fabricated. The average relative density of fabricated samples reaches up to 99.96%. The ultimate tensile strength of Ti-6Al-4V sample was 923~1174 MPa, comparable to that of wrought parts. Some Ti-6Al-4V parts fabricated by EBSM are showed in Fig. 10.Slightly sintered powderAfter cleaningFig. 10. Ti-6Al-4V parts fabricated by EBSM3.5. Dual-Materials ProcessingA Ti-6Al-4V and TiAl dual-material structure is proposed and planned to be fabricated on the system developed in this paper. Ti-6Al-4V has excellent strength and plasticity at room temperature, while TiAl alloy has excellent performance at high temperature, yet very brittle. Part with Ti-6Al-4V at the cold end, TiAl at the hot end, and graded material at the middle may have potential application prospect. As illustrated in Fig. 11, at the bottom layers material with 100% Ti-6Al-4V and 0% TiAl is deposited.

As illustrated in Fig. 11, at the bottom layers material with 100% Ti-6Al-4V and 0% TiAl is deposited. Then the proportion of TiAl increases from 0% to 100% layer by layer, finally turning into 100% TiAl at the top layers. This structure withdual-performance can be used to build the TiAl impeller with a Ti-6Al-4V rabbet. Compared with the welding interface, the graded material transition would be more controllable and programmable.However, the dual-materials processing brings challenges to the powder recycle. As the graded parts are fabricated in a single machine run, the two powders are mixed in the building tanks. As a result, the un-used powder is difficult to be recycled and is sacrificed to obtain the graded material.100% TiAl80% TiAl, 20% Ti-6Al-4V60% TiAl, 40% Ti-6Al-4V40% TiAl, 60% Ti-6Al-4V20% TiAl, 80% Ti-6Al-4V100% Ti-6Al-4VFig. 11. Ti-6Al-4V and TiAl dual-material structureConclusions1.

11. Ti-6Al-4V and TiAl dual-material structureConclusions1. The vibration driven powder supplier developed in this study has a high compatibility to various powders. A stable supplying rate and a supplying accuracy less than 7.5% were obtained with the supplier.2. Two powders can be supplied individually to obtain a mixture with tailored proportion. Experimental results show that the mixture is homogenous and the actual proportion is close to the desired value.3. The powder spreading device based on dual inclined combs can significantly reducethe stress of comb tooth while scraping over a bump and thus has longer lifetime.4. The EBSM system provides two exchangeable building tanks with sizes of 100 × 100× 100 mm3 and 250 × 250 × 250 mm3.5.

The EBSM system provides two exchangeable building tanks with sizes of 100 × 100× 100 mm3 and 250 × 250 × 250 mm3.5. The novel EBSM system is capable of building parts with single material and has apotential of dual-material processing.AcknowledgementsThe authors would like to acknowledge the funding of 2013 Beijing Science andTechnology Development Project, No. D13110400300000 and D131100003013002.References1. Cormier, D., Harrysson, O. L. A., Mahale, T., West II, H. A., 2008, “Freeform Fabrication of Titanium Aluminide via Electron Beam Melting Using Prealloyed and Blended Powders”, Advanced Material Science and Engineering, 2007, pp. 6822-6825.2. Facchini, L., Magalini, E., Robotti, P., Molinari, A., 2009, “Microstructure and mechanical properties of Ti-6Al-4V produced by electron beam melting of pre-alloyed powders”, RapidPrototyping Journal, 15, pp. 171-178.3. Murr, L. E., Martinez, E., Pan, X. M., Gaytan, S. M., Castro, J. A., Terrazas, C.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDEVELOPMENT OF NOVEL HIGH TEMPERATURE LASER POWDER BED FUSIONSYSTEM FOR THE PROCESSING OF CRACK-SUSCEPTIBLE ALLOYSL. Caprio*, G. Chiari*, A.G. Demir*, B. Previtali*Department of Mechanical Engineering, Politecnico di Milano, Via La Masa 1, 20156, Milano, ItalyAbstractIn the industrial panorama, Laser Powder Bed Fusion (LPBF) systems enable for the near net shaping of metal powders into complex geometries with unique design features. This makes the technology appealing for many industrial applications, which require high performance materials combined with lightweight design or conformal cooling channels. However, many of the alloys that would be ideal for the realisation of these functional components are classified as difficultly weldable due to their cracking sensitivity.

However, many of the alloys that would be ideal for the realisation of these functional components are classified as difficultly weldable due to their cracking sensitivity. Currently, industrial SLM systems employ baseplate preheating to minimise these effects although this solution is limitedly effective along the build direction and often does not achieve high enough temperatures for the realisation of crack-free specimen. In this work, the design and implementation of a novel inductive high temperature LPBF system is presented.

In this work, the design and implementation of a novel inductive high temperature LPBF system is presented. Furthermore, preliminary results regarding depositions of Titanium Aluminide alloy with and without preheating are reported, showing the potential of the solution developed.Keywords: Laser Powder Bed Fusion, High Temperature, Preheating, TiAlIntroductionLaser Powder Bed Fusion (LPBF), also known as Selective Laser Melting (SLM), employs a highly energetic laser beam to selectively melt a layer of metallic powder feedstock which is laid upon a substrate material. The process is thus repeated layer-by-layer as typically done in other Additive Manufacturing (AM) processes. In order to obtain stable processing conditions and high part density, full melting of the powder bed is required.

In order to obtain stable processing conditions and high part density, full melting of the powder bed is required. Hence, for the processing of metallic powders, LPBF requires the use of a beam with elevated energy density which has been enabled by the development of high brilliance fiber laser sources in recent years [1]. Although, the Laser Powder Bed Fusion technology has been witnessing an increasing interest on behalf of the industrial world, its applicability is mainly restricted to the processing of weldable alloys. The main reason behind the limitation in the range of materials processable is linked to the elevated thermal gradients and cooling rates that are experienced locally during the fusion process and are intrinsic to the technology due to the concentrated energy beam that is required to melt the powder bed. Cooling velocities greater than 106 K/s have been measured experimentally by Scipioni et al. [2] while Shi et al.

Cooling velocities greater than 106 K/s have been measured experimentally by Scipioni et al. [2] while Shi et al. estimate values of thermal gradients above 104 K/mm [3]. As clearly explained by Hagedorn et al. these values are critical when processing metallic powders, due to the elevated stress concentrations generated as a consequence [4].Powder bed preheating has thus been extensively employed since the onset of the laser powder bed fusion technology [5]. Initially, its role was that of lowering the energy density required to achieve full melting of the feedstock material but with the introduction of high brilliance fiber lasers the main scope became the minimisation of thermal tensions which form locally. As a matter of fact, all of the industrial systems are equipped with preheating systems that allow the reduction of thermal stresses in the first layers of the building process.

As a matter of fact, all of the industrial systems are equipped with preheating systems that allow the reduction of thermal stresses in the first layers of the building process. In the initial depositions of the AM process, the reduction of thermal stresses is essential due to the presence of the substrate base which acts as a heat sink and thus causes the formation of higher thermal gradients and cooling rates.Different techniques of substrate preheating have been applied to the laser powder bed fusion technology, namely IR heaters [5], preheating or remelting strategies through fast scanning of the laser beam [6,7], use of a second defocused laser beam [8], substrate resistive heating [9,10] or baseplate induction circuit [4,11,12]. The greater part of industrial selective laser melting systems employs resistive heaters [10], which typically allow2275preheating of the baseplate up to 200°C while the most recently developed systems enable to achieve up to 500°C.

Although these systems are particularly effective in reducing local stresses near the base plate, they lose their effectiveness when a full-scale component must be manufactured since their effect remains limited to the lower parts of the build (due to the inherent nature of the process which relies on the lowering of the base plate). A similar preheating concept was developed by Hagedorn et al. which on the other hand employed a pancake induction heating circuit capable of achieving up to 1000 °C [11,12]. IR heaters have been applied for the preheating of the top layer of the powder bed, however these systems are not energetically efficient [5,13].Alternatively, the use of a second defocused laser beam may be effective in reducing residual stresses as estimated by Aggarangsi et al. [14] and demonstrated experimentally Wilkes et al.[8].

[14] and demonstrated experimentally Wilkes et al.[8]. Although, this approach is effective when processing non-metallic materials, it implies a higher capital cost of the AM machine since a second optical set up and laser source must be equipped for this solution. Preheating or remelting strategies employing the same process laser can be effective in reducing porosity as demonstrated by Demir et al.[7] and Aboulkhair et al.[6]. However, these imply a notable increase in terms of production time due to the double scanning of the powder bed required. A competing technology which relies on a similar preheating concept is Electron Beam Melting (EBM) which is capable of achieving powder bed temperatures in the order of 1000°C whilst maintaining high build rates due to the higher scanning velocities enabled by the magnetic coil deflection system [15]. For this reason, the EBM technology is capable of processing γ-TiAl alloys.

For this reason, the EBM technology is capable of processing γ-TiAl alloys. However, a downside of the EBM is that the minimum feature achievable is limited by the melt pool dimension which is typically greater in comparison to laser powder bed fusion systems [16].Taking into consideration the available technologies and the need for uniform high preheating temperatures for the processing of crack susceptible alloys, the aim of the present work was the development of a novel high temperature laser powder bed fusion (HT-LPBF) system, namely project Grisù. The preheating system chosen for this purpose was an induction coil with a different geometry with respect to that implemented by Hagedorn et al.[4]. The EBM technology or IR heating systems are capable of heating the last few layers of the powder bed. On the other hand, baseplate inductive or resistive heating systems are effective only for low build heights.

On the other hand, baseplate inductive or resistive heating systems are effective only for low build heights. The solution developed in the current research enables a constant preheating temperature throughout of the whole build. Instead of employing a pancake coil, the novel HT-LPBF was designed with a helical coil which contains the substrate base as it is lowered during the building process. The use of this innovative system architecture might not only open the road for the processing of crack-susceptible alloys but also for the realisation of in-situ heat treatments during the freeform fabrication process.The scope of this article will therefore present this new conceptual AM system and is structured as follows: the design of project Grisù is presented and the testing of the induction system to achieve preheating temperatures above 800°C is reported.

Eventually a demonstrator build using a TiAl alloy prone to crack formation is realised with and without the use of the preheating system. Results show that the prototypal system designed is effective in suppressing crack formation and acts as a proof of concept of the novel HT-LPBF system architecture.2276Design of a novel high temperature LPBF systemDesign of the inductive preheating systemProject Grisù was born with the aim of developing an innovative additive manufacturing system whilst maintain the advantages of an open-ended laboratory set-up. For this reason, the prototypal system was designed to operate with low quantities of powder, similarly to its predecessor, project Powderful which consisted in a basic automated powder deposition system[17]. The novel conceptual design of the preheating system compared to the classical resistive baseplate heating is schematically shown in Figure 1.Figure 1.

The novel conceptual design of the preheating system compared to the classical resistive baseplate heating is schematically shown in Figure 1.Figure 1. Conceptual layout design of (a) EBM with pre-scanning heating (b)LPBF with standard resistive preheating system and (c) novel LPBF with inductive heating systemA critical aspect of the design of this prototype was the choice of materials for the different components. It was essential that the most stressed components of the system could sustain temperatures up to 800 °C whilst not being influenced by the strong magnetic field generated by the induction coil. For this reason, quartz glass was selected for the cylinder containing moveable part of the z-axis. The piston was realised in Cerastil (Formenti, Milano, Italy) a poly-silicate material capable of withstanding up to 1100°C.

The piston was realised in Cerastil (Formenti, Milano, Italy) a poly-silicate material capable of withstanding up to 1100°C. Accordingly, the gasket of the piston which prevents powder infiltration in the Z-axis was realised in Paperseal (Formenti, Milano, Italy) a mineral fibre material. The fixed plane of the powder bed was designed to be in ceramic glass whilst the helical coil for the preheating system was of pure copper.Within the piston, a small steel cylinder was located which inductively heats up during the printing process and contributes to the preheating of the metallic substrate by means of conduction. In case the prototypal system is employed to study the processing of non-metallic materials which must thus be deposited upon non-metallic substrates, the steel cylinder will act as the heat source for the preheating system. In the present work, for the testing of project Grisù the substrate material chosen was Ti6Al4V.

In the present work, for the testing of project Grisù the substrate material chosen was Ti6Al4V. Titanium alloys although they are light non- ferrous materials tend to behave similarly to steel when heated inductively (since they possess analogous values of electrical resistivity and thermal conductivity) [18]. Hence, the whole substrate base can be expected to be effectively heated up by means of the inductive preheating system. On the other hand, during the printing process the induction field will also have an influence over the powder bed. Although, the effect in terms of inductive preheating on metallic powders is unclear in literature and will be the focus of future studies, the system will still be effective in preheating the solidly deposited material.The “heart” of project Grisù was designed with three water heat exchangers in order to contain the elevated temperatures being generated from the induction heating system.

An adequate cooling system is fundamental to avoid distortions of other mechanical parts and in particular the accuracy of the Z-axis movement system. The first heat exchanger is the induction coil itself which is composed of copper tubes where water is forced to flow. This acts not only as a barrier for the heat flow with respect to the rest of the machine but also plays an important role in maintaining the copper tubes cooled. A second flat water-cooled heat exchanger is located in a fixed position beneath the heart of the induction system. The third water-cooled component translates along the Z-axis as it is lowered during the build job and is responsible of maintaining at a low temperature the motor transmission.2277Temperature sensors were placed on these components in order to monitor their temperature and avoid them from achieving excessive values.

The components and their respective materials are listed in Figure 2, alongside with a side view of the design of the preheating system and Z-axis. Tmax refers to the maximum operational temperature of the different components (depending on their material). These were indicated within the specifications given by the part producers or providers.

Tmax refers to the maximum operational temperature of the different components (depending on their material). These were indicated within the specifications given by the part producers or providers. Although, relatively high values of Tmax are indicated for the mechanical components the heat exchangers design is supposed to maintain these values below 100 °C.#ComponentMaterialT max [°C]System design1PistonCemestil11002GasketPaperseal11003PlaneCeramic glass8004CylinderQuartz glass12505Induction coilPure Cu2506Fixed heat exchangerAl-alloy4007 Moving heat exchangerAl-alloy4008-9 Nut and screwCarbon steel50010Piston-Heat exchanger connection Carbon steel50011 Rebored barCarbon steel50012 Linear bearingSteel50013 Motor transmissionElastomer7014Stepper motorMixed7015Steel insertCarbon steel100016 Thermocouple positionFigure 2.

List of materials and their maximum operational temperature T max and design of the preheating systemThe final design and realisation of project Grisù is shown in Figure 3. The recoater blade was realised in stainless steel in order to withstand the elevated temperatures of the powder bed. The system is equipped also with a fan which provides a gas flow above the build area. The whole prototypal system can be placed within an inert gas chamber (analogously to the previous prototypal system developed [17]). Before each processing the powder bed and activating the preheating system, an inertization procedure of the closed chamber was conducted, applying a vacuum down to 50 mBar and then filling the process chamber with high purity Ar gas. The whole mechanical control of the system is conducted with a self-developed LabVIEW code (National Instruments, Austin, TX, USA).2278Figure 3.

The whole mechanical control of the system is conducted with a self-developed LabVIEW code (National Instruments, Austin, TX, USA).2278Figure 3. Project Grisù (a) CAD design (b) assembled prototype system (c) final set-upLaser optical delivery systemFor the present research, a 1000 W single-mode fibre laser source (n-Light, Vancouver, WA, USA) was employed as a light source to process the powder bed. The beam path was controlled with a Smart Scan SH30G- XY2 scanner head (Smart Move GmbH, Garching bei München, Germany), whereby the laser emission was collimated with a 75 mm lens and focused onto the substrate using a 420 mm F-theta lens. The beam diameter on the focal plane was determined from theoretical calculations as d0= 78 μm. Overall, the principal features in terms of optical and mechanical parameters of project Grisù are reported in Table 1.Table 1. Characteristic parameters of project Grisù, the HT-LPBF systemParameter Laser source Laser max.

Characteristic parameters of project Grisù, the HT-LPBF systemParameter Laser source Laser max. power, Pmax Beam quality factor, M2 Max. scan speed, v scan,max Spot diameter, d0 Layer thickness, z Process atmosphere Max preheating temperature, Tpreheat,max Build plate diameter, ϕValue nLight Alta Prime 1 kW 1.19 5 m/s 78 µm 30 to 100 µm Ar; N2 >800°C 38 mmTesting of the high temperature preheating systemIn order to verify the functionality of the high temperature preheating system a measurement chain to keep track of the temperature variations within the powder bed had to be designed. A K-type thermocouple was employed for this scope (TERSID, Milano, Italy) which is indicated for use in these processing conditions by the producer. The measurement point was selected just beneath the substrate position such that the actual temperature of the baseplate could be measured (as indicated in Figure 2).

The measurement point was selected just beneath the substrate position such that the actual temperature of the baseplate could be measured (as indicated in Figure 2). Although this configuration does not allow to2279monitor the temperature of the whole system, given the reduced dimensions of the powder bed it might be taken as representative of the effective thermal history of the build. An alternative solution for the temperature control could be the use of a pyrometer but this equipment has higher capital cost and was therefore not taken into consideration as an option during the realisation of a prototypal system whose main objective is the proof of concept of a novel preheating system. Future developments of the system could also attempt at using more thermocouples to monitor the temperature along the build direction.Figure 4.

Future developments of the system could also attempt at using more thermocouples to monitor the temperature along the build direction.Figure 4. (a) Incandescent Ti6Al4V substrate during the testing of the high temperature induction preheating system and (b) HT-LPBF of a γ-TiAl alloy on project GrisùThe control of the powder bed temperature was obtained through an ON/OFF solution, whereby the magnetic field was switched on and off. A Zero Voltage Switching (ZVS) circuit was employed to generate the current that was fed to the helical coil to generate the magnetic field. A 48V-20A DC power supply (Mean Well, New Taipei, Taiwan) was employed to power the ZVS circuit. A switch to achieve the ON/OFF control was interposed between the power supply and the ZVS circuit in order to avoid the transient behaviour of the power supply.

A switch to achieve the ON/OFF control was interposed between the power supply and the ZVS circuit in order to avoid the transient behaviour of the power supply. The resonating frequency of the inductive circuit was estimated as 100 kHz.In order to verify the functionality, the inductive preheating system was tested initially up to 200°C, 500°C and 700°C. The maximum temperature achievable by the inductive preheating system was in excess of 840°C so the final temperature testing condition which simulated printing conditions and times was conducted at 800°C. In all tested conditions the power input to the induction system corresponded to maximum power output of the DC supply which is 960 W. Figure 4 (a) shows an incandescent Ti6Al4V substrate during one of the testing phases while Figure 4 (b) depicts the powder bed at 800°C during the deposition of a γ-TiAl alloy.

Figure 4 (a) shows an incandescent Ti6Al4V substrate during one of the testing phases while Figure 4 (b) depicts the powder bed at 800°C during the deposition of a γ-TiAl alloy. Each of the preheating temperature tests consisted of four phases: heating, maintenance, controlled cooling and convective cooling. The heating phase must avoid excessive thermal shock on the system components but in any case results less critical since the component of a crack susceptible alloy is yet to be built. The time to achieve the operating temperature value of the system was set at 30 minutes. The maintenance phase at 800°C was tested for a 450 minute duration whereby the temperature oscillated with a ±10°C margin of the target temperature. Finally, the controlled cooling phase is shown, whereby the system was tested to maintain a cooling ratio of 5°C/min (always controlled using the ON/OFF control logic previously reported).

Finally, the controlled cooling phase is shown, whereby the system was tested to maintain a cooling ratio of 5°C/min (always controlled using the ON/OFF control logic previously reported). Finally, a convective cooling of the system was experimented in order to view the behaviour of the system. This was arbitrarily tested from 450 °C downwards.In Figure 5, the temperatures recorded by the thermocouple system are reported during the different phases of the testing at 800 °C (selected time instants are reported). Although, the ON/OFF control introduces an error into the powder bed temperature with respect to the target temperature, it is also true that this simple control logic system is functional for the aim of the present work. The largest errors with this control logic may be witnessed at temperatures below 500°C since even if the induction circuit is switched on for a very brief period it rapidly raises the temperature of the system.

At higher temperatures, however the error reduces and may be considered as acceptable.2280Figure 5. Testing of the inductive preheating system up to 800°C. Powder bed temperature measured by thermocouple, T bed in blue whilst target temperature, T target when present in orange. Different phases are shown: (a) heating (b) maintenance (c) controlled cooling (d) convective cooling.Testing of the high temperature laser powder bed fusion systemMaterialIn order to verify the functionality of the novel HT-LPBF, testing of the system was conducted with a γ- TiAl alloy. As can be evinced by the literature on the laser melting of titanium aluminide alloys, this material is prone to crack formation due to the elevated temperature gradients and cooling rates of the process as well as the ductile-brittle transition of the alloy (which tends to occur around 750°C as stated by Vilaro et al.) [19]. Löber et al.

Löber et al. witnessed the presence of cracking already during single track deposition of a γ-TiAl alloy [20]. Similarly, Doubenskaia et al. with a preheating up to 450°C witnessed cracks during the realisation of cubic specimen of Ti–48Al–2Cr–2Nb [21].Figure 6. Scanning Electron Microscope images of Ti-48Al-2Cr-2Nb gas atomised powder low magnification (left) and high magnification (right).Since Project Grisù was successful in maintaining temperatures of the powder bed around 800°C it should be apt for the processing of γ-TiAl alloys. Specifically, Ti-48Al-2Cr-2Nb gas atomized powder (LPW Technology, Runcorn, UK) was tested on the novel HT-LPBF system. The chemical composition of the feedstock material as declared by the producer is reported in Table 2 while the size distribution parameters are: d10=16 μm, d90=45μm. Scanning electron microscope images of the γ-TiAl powder are visible in Figure 6.Table 2.

Scanning electron microscope images of the γ-TiAl powder are visible in Figure 6.Table 2. Chemical composition of Ti-48Al-2Cr-2Nb declared by the producerElement Ti Wt %Si Bal. 34.4 2.58 4.74 0.04 0.006 0.007 0.12 0.006 0.001AlCrNbFeCONH2281Measurement apparatusQualitative evaluation of the top surface and of the polished cross-sections of the specimen deposited was done through optical microscopy images, acquired with a 2.5X objective (Quick Vision ELF from Mitutoyo, Kawasaki, Japan).Experimental conditionsLaser powder bed fusion of the Ti-48Al-2Cr-2Nb powder was investigated into two different experimental conditions (i.e. at ambient temperature and at T preheat=800 °C). Preliminary experimentation led to the choice of different laser energy density delivered to the powder bed, respectively corresponding to 357 J/mm3 for the print at ambient temperature and 119 J/mm3 in the case of Tpreheat=800 °C.

For both experimental conditions, laser power was fixed at 150 W, scan speed at 200 mm/s and hatch distance at 70 μm. Energy density variations were controlled processing the powder bed with different the layer thicknesses whilst the preheating contribution was not taken into account since its effect is two orders of magnitude lower (i.e. 2 J/mm3). In the case of the laser deposition with high preheating temperature, once the build was completed a controlled cooling rate, ΔṪ =5 °C/min was applied as indicated by Gussone et al.[12]. Within both configurations, the inertization gas employed was high purity Argon and the substrate material was made of Ti6Al4V. 4x4 mm2 square hatched laser trajectories were projected on to the powder bed in order to realise cubic specimen layer-by-layer.ResultsBoth experimental conditions, with and without the high temperature preheating of the powder bed it was possible to deposit material onto the Ti6Al4V substrate.

However, in the case of the ambient temperature printing of the γ-TiAl alloy the process had to be interrupted after 10 layers of material had been deposited due to powder bed instabilities and protruding parts which impeded the correct deposition of successive layers. Contrarily, the high temperature preheating at 800°C was capable of maintaining a stable powder bed and 50 layers were successfully deposited. These instabilities can be clearly viewed from the top surface finish of the specimen realised, shown in Figure 7.Figure 7. Top view of specimen acquired with optical microscope (a) no preheating (b) with preheating at 800°CFurthermore, it is possible to denote the extensive presence of cracks in the specimen realised at ambient temperature, clearly evident in the top view of Figure 7 (a) and also confirmed by the cross-section of the specimen shown in Figure 8 (a)).

The crack density for the specimen realised was measured to be from the metallographic2282analysis as 4.23 mm/mm2 which indicates that without the high temperature preheating system, it is impossible to realise components in γ-TiAl since the presence of defects highly compromises the structural integrity of the components realised. During the present experimentation cracking occurred predominantly perpendicularly to the scanning direction of the laser beam. This is in agreement with the findings of Löber et al. whom identified the cracking always perpendicularly to the scanning direction during the single track deposition of a γ-TiAl alloy [20]. Furthermore, the formation of cracks transverse to the advancement of the heat source has been reported for process to laser powder bed fusion (i.e. electron and laser beam welding) [22,23].Figure 8.

electron and laser beam welding) [22,23].Figure 8. Cross section along build direction of specimen acquired with optical microscope (a) no preheating (b) with preheating at 800 °CThe reason for the formation of such defects can be related to the ductile to brittle transition which occurs around 750 °C and with the local strong thermal stresses generated by the process. γ-TiAl alloys are known to be susceptible to this phenomenon, often referred to as “cold cracking” [24]. The use of local preheating to minimise thermal stresses has been identified as a solution for the defect-free processing of Titanium-Aluminides [22–24].

The use of local preheating to minimise thermal stresses has been identified as a solution for the defect-free processing of Titanium-Aluminides [22–24]. The absence of cracks in the cross-section of the specimen deposited with high temperature preheating confirms results coming from previous literature studies and shows promising results for the development of industrial systems capable of realising fully functional γ-TiAl components.Conclusions and future developmentsThe main objective of the present work, namely the realisation of a novel high-temperature laser powder bed fusion system, was achieved. Project Grisù was successfully tested to preheat the powder bed up to 800 °C. A demonstrator print of Ti-48Al-2Cr-2Nb was conducted both with and without preheating in order to verify the functionality of the system in processing crack-susceptible alloys.

A demonstrator print of Ti-48Al-2Cr-2Nb was conducted both with and without preheating in order to verify the functionality of the system in processing crack-susceptible alloys. The temperature control implemented was effective in maintaining the temperature within the desired range and in controlling the cooling rate such to prevent crack formation of the deposited samples. Specimen processed without the use of the inductive system presented severe cracks. The defects developed perpendicularly to the scanning direction of the laser beam and are related to the cold cracking formation mechanism. On the other hand, components realised at Tpreheat=800°C, did not show the presence of cracks under optical microscope inspection.

On the other hand, components realised at Tpreheat=800°C, did not show the presence of cracks under optical microscope inspection. Further characterisation of the deposited specimen is required in order to confirm the absence of cracks although preliminary analyses are promising.The potential of the system developed principally relies upon its ability of uniformly heating up the powder bed and the component being realised. This opens up the barrier for in-situ heat treatment of the component during its freeform fabrication which is not possible within the technology currently available. Nonetheless, it must be taken into consideration that project Grisù is a prototypal system with reduced powder bed dimensions and this greatly simplifies the management of the heat input. In the case a full-scale industrial system should be designed a much greater attention should be directed towards the containment and distribution of temperature within the system.

In the case a full-scale industrial system should be designed a much greater attention should be directed towards the containment and distribution of temperature within the system. Furthermore, open questions remain with regards to technological limitations of the process and mechanical properties of components realised as well as the effect of the inductive field on the process dynamics. This technological development has the potential of enabling a whole new perspective in the freeform fabrication2283of materials which to the present moment could only be processed with conventional technologies or complex manufacturing routes yet many unanswered queries remain open for research.AcknowledgmentsThe authors wish to express their gratitude to Dr. Eligio Grossi for his contribution in the design of the prototype system. Acknowledgments go also to Optoprim for providing the controller system and to nLight for kindly providing the laser source.

Eligio Grossi for his contribution in the design of the prototype system. Acknowledgments go also to Optoprim for providing the controller system and to nLight for kindly providing the laser source. This work was supported by European Union, Repubblica Italiana, Regione Lombardia and FESR through the project MADE4LO under the call "POR FESR 2014-2020 ASSE I - AZIONE I.1.B.1.3”.References[1] Demir AG, Colombo P, Previtali B. From pulsed to continuous wave emission in SLM with contemporary fiber laser sources: effect of temporal and spatial pulse overlap in part quality. Int J Adv Manuf Technol. (2017)91(5–8):2701–14. Scipioni Bertoli U, Guss G, Wu S, Matthews MJ, Schoenung JM. In-situ characterization of laser-powder interaction and cooling rates through high-speed imaging of powder bed fusion additive manufacturing. Mater Des. (2017)135:385–96. Shi Q, Gu D, Xia M, Cao S, Rong T.

Investigation of remelting and preheating in SLM of 18Ni300 maraging steel as corrective and preventive measures for porosity reduction. Int J Adv Manuf Technol. (2017)93(5–8):2697– 709.[8] Wilkes J, Hagedorn Y, Meiners W, Wissenbach K. Additive manufacturing of ZrO 2 ‐Al 2 O 3 ceramiccomponents by selective laser melting. Rapid Prototyp J. (2013)19(1):51–7.[9] Kempen K, Vrancken B, Buls S, Thijs L, Van Humbeeck J, Kruth J-P. Selective Laser Melting of Crack- Free High Density M2 High Speed Steel Parts by Baseplate Preheating. J Manuf Sci Eng. (2014)136(6):061026.[10] Ali H, Ma L, Ghadbeigi H, Mumtaz K. In-situ residual stress reduction, martensitic decomposition and mechanical properties enhancement through high temperature powder bed pre-heating of Selective Laser Melted Ti6Al4V. Mater Sci Eng A. (2017)695(February):211–20.[11] Gussone J, Garces G, Haubrich J, Stark A, Hagedorn YC, Schell N, et al. Microstructure stability of γ-TiAlproduced by selective laser melting.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDevelopment of Novel Tapered Pin Fin Geometries for Additive Manufacturing of Compact Heat ExchangersJulien Cohen1, David L. Bourell1,21Mechanical Engineering Department 2The Center for Materials Science and Engineering The University of Texas at Austin Contact: jhco@utexas.eduAbstractPin fin arrays are widely used to enhance forced convection heat transfer across various industries, finding application in turbine blade trailing edges, electronics cooling, and broadly for compact heat exchange. Fin shape greatly affects flow separation and turbulence generation, and optimizing performance relies on this balance between increased heat transfer and increased pressure loss along the array.

Fin shape greatly affects flow separation and turbulence generation, and optimizing performance relies on this balance between increased heat transfer and increased pressure loss along the array. Straight circular and elliptical fins are well-characterized in the literature, and there exist a scant few studies on tapered configurations with conventional cross- sections. Recent works have investigated straight pin fins with more complex shapes. Tapered, complex fin geometries represent an avenue for overall performance gains, but manufacturing them is difficult and time-consuming using traditional machining processes. The unique capabilities of additive manufacturing now allow their economical fabrication in an increasing number of fully-dense engineering materials.

The unique capabilities of additive manufacturing now allow their economical fabrication in an increasing number of fully-dense engineering materials. This work compares 21 pin fin arrays of varying fin cross-section, taper angle, taper profile, and array patterns using experimental and computational methods.IntroductionPin fin arrays are widely used to enhance forced convection heat transfer across various industries, finding application in turbine blade trailing edges, electronics heat sinks, compact heat exchangers, and broadly wherever high heat flux removal is required in compact volumes. The use of pin fins increases heat transfer through two main mechanisms: increasing the wetted surface area over which convective heat transfer can occur, and promoting turbulent flow in the inter-fin region of the array [1]–[4]. This turbulence is generated mainly by interrupting boundary layers on the channel walls, and inducing vorticity at the base and in the wake of the fins [5].

This turbulence is generated mainly by interrupting boundary layers on the channel walls, and inducing vorticity at the base and in the wake of the fins [5]. Wake flow regimes within these arrays are characterized by complex transitions and strongly three-dimensional flow effects [6].However, increases in flow obstruction and turbulence are directly related with an increase in pressure loss along the fin array. For many applications where cooling fluid flow is derived from the system itself, specifically in modern gas engine turbine blades used in the aerospace industry, it is desired to minimize this pressure loss to increase overall system efficiency [7]. Even in systems using discrete pumps or fans to provide coolant flow, reducing pressure loss means the system draws less power, is more efficient, and is more environmentally friendly, to produce an equivalent level of heat transfer.

Therefore, attaining the highest level of performance from a pin fin array of this type is a nonlinear optimization problem which must holistically balance fin cross-section, taper profile, array spacing, channel geometry, surface2314features, and fluid flow characteristics. The history of academic research in this field shows the consistent movement towards higher streamwise aspect ratio, curved, variable profile pin fins.Conventional circular pin fins are well-characterized in the literature [3], [4], [6], [7]. Tapered pin fins with circular cross-section have received very little public scrutiny, but are investigated in [8], [9]. More recent works have investigated straight pin fins with complex geometries including elliptical [1], [2], [5], [10]–[15], square [5], [16], [17], triangular [5], [6], airfoil derived [5], [13], and teardrop cross-sections [5], [11], [13], [18].

To date, this author has found no example of investigations on tapered pin fins with novel or non-conventional cross- sections.This work explores the potential performance enhancements of using tapered, complex cross-section pin fin arrays for forced convection heat exchange. Twenty-one array designs, separated into eight geometry families, are investigated using experimental and computation means. Comparisons are made between the members of each geometry family, as well as between families, to illuminate effective means for increasing overall heat exchanger performance. Non-dimensional metrics are defined for heat transfer enhancement, pressure loss, and overall performance to quantitatively compare disparate array configurations. Topics for future work and potential methods for investigation are suggested.MethodologyCross-sectional geometries were defined for teardrop shaped pin fins to be used in this work.

Topics for future work and potential methods for investigation are suggested.MethodologyCross-sectional geometries were defined for teardrop shaped pin fins to be used in this work. Each pin consists of a circular section and a tail section defined by two bilaterally symmetric, tangent extensions. Teardrop tail length L and radius R are defined as parameters, as well as a span-wise and stream-wise pitch for the pin pattern, yn and xn (Figure 1).Figure 1: Cross-section of a single pin, and cross-section of three pins within the exchanger pattern.Figure 2: Area normalized teardrop pin geometries with aspect ratios varying from N = 0.1 to N = 0.9.2315An aspect ratio, N, is defined as the ratio of radius and tail length.

Using the geometry of the pin, a formula is defined for the radius dependent only on N and the cross-sectional area of the pin, A.𝑁 =𝑅𝐿(1)𝑅 = √𝐴 𝜋−𝑐𝑜𝑠−1 𝑁+√𝑁−2−1(2)In choosing the form of these equations, it was determined that normalizing by area was of more utility than normalizing by perimeter. The area of each pin more directly correlates with the time and expense required to laser sinter it, as well as the final mass, which is important for performance-driven applications. By holding A constant and defining a set of N values from 0.1 to 0.9, a range of R and L values are calculated. These define the range of potential pin geometries for analysis, shown in Figure 2. The geometry is quantified based on normalization to a circle of radius 3.75 mm.Three pin geometries were chosen to be tested from this range, to represent the respective extreme ends of the spectrum as well as an intermediate case.

0.2AR, 0.6AR, and 0.8AR are used throughout this work to denote these chosen aspect ratios. Each of the three define a geometry family whose members all have the same initial cross-section at the base of the fin. Five more families are defined for testing, summarized in Table 1 and discussed in the following subsections.Name 0.2ARMembers Description3High aspect-ratio teardrop shaped fins with a length-radius ratio of 0.2, skinny and long. Straight, low taper, and high taper versions. Medium aspect-ratio teardrop shaped fins with a length-radius ratio of 0.6. Straight, low taper, and high taper versions. Low aspect-ratio teardrop shaped fins with a length-radius ratio of 0.8, short and stubby. Straight, low taper, and high taper versions. Based on the 0.6AR design, these pins decrease in diameter along the streamwise length of the array. Straight and low taper versions. Based on the 0.6AR design, these arrays feature nonlinear tapers along the height of each fin.

Straight and low taper versions. Based on the 0.6AR design, these arrays feature nonlinear tapers along the height of each fin. Convex and concave versions. The fins on these arrays are Standard Elliptical Fins (SEF), in a straight, low taper, and high taper version. Based on the 0.6AR design, these fins are swept backwards along the height of each fin. Straight and high taper versions. Based on the 0.6AR design, these use a pattern which repeats every three rows, rather than every two. Straight and high taper versions.0.6AR30.8AR3Diameter change Nonlinear taper SEF223Swept back Triple row Table 1: Summary of tested geometry families.22316Four of these families are based on the 0.6AR geometry, with novel design features varying the fin shape along the vertical or streamwise axis. The fifth family is based on the Standard Elliptical Fin (SEF), which previous work has shown is more efficient than circular fins in a non-tapered configuration [1], [13], [14].

The fifth family is based on the Standard Elliptical Fin (SEF), which previous work has shown is more efficient than circular fins in a non-tapered configuration [1], [13], [14]. Each family contains members with varying tapers applied to the fin extrusion, defined by a draft angle relative to the vertical axis. “Straight” members have no taper, effectively a draft angle of 0°. “Low Taper” members have a draft angle of 2-3°, and “High Taper” members have a draft angle of 4-5°, dependent on geometry limitations. Each fin array is patterned with pitches of xn = 10 mm and yn = 20 mm, with 10.8 mm of flow bypass on each side of the array and 30 mm tall fins. The base of the arrays is 6 x 4 x 0.1 in. Every array pattern repeats every two rows with an offset of 10 mm between, except for the two triple repeating designs, which repeat every three rows with an offset of 6.67 mm.1 2 34 5 6Figure 3: 1: 0.2AR individual fin cross sections: (a) straight, (b) low taper, (c) high taper.

2: 0.6AR individual fin details: (a) straight, (b) low taper, (c) high taper. 3: 0.8AR individual fin details: (a) straight, (b) low taper, (c) high taper. 4: Nonlinear tapered fin array geometries: (a) concave, (b) convex. 5: SEF individual fin details: (a) straight, (b) low taper, (c) high taper. 6: Swept back individual fin details: (a) straight, (b) high taper.Experimental TestingEach of the pin fin arrays were fabricated using Nylon 12 PA powder supplied by Stratasys Direct Manufacturing [19] in a 3D Systems Sinterstation HiQ+HS machine. The powder stock was a mixture of fresh powder, overflow, and part cake, at approximately a 30/40/20 percent mass ratio.2317A prototype wind tunnel system was created for testing the real-world pressure drop over the various pin fin arrays.

This system was designed for a modular, open-loop configuration utilizing a centrifugal blower fan to produce inlet fluid flow with Reynold’s number 0-104, allowing drop-in testing of any fin array within the size limitations. Two differential pressure transducers were configured to measure fluid velocity using a pitot-static probe upstream of the test section, and pressure drop over the test section using pressure ports. Data from these transducers were fed through a USB data acquisition (DAQ) module into purpose built Labview software for real-time test monitoring and data logging. Figure 4 shows an overview of the wind tunnel system architecture.Figure 4: Diagram of wind tunnel system architecture.A Dayton model 3HMJ1 centrifugal blower (1) uses a squirrel cage rotor to draw in air and accelerate it through the system. The air travels through an expansion section (2) into the settling chamber (3), which allows large-scale turbulent fluctuations to decay [20].

The air travels through an expansion section (2) into the settling chamber (3), which allows large-scale turbulent fluctuations to decay [20]. The air passes through a honeycomb filter (4) which aligns the flow parallel to the test section and minimizes lateral variations in the mean and fluctuating velocity [21]. The air then flows through a contraction section (5) into the test section (6), where it interacts with the fin array (9), secured with an airtight seal by the array housing (10). This seal is effected by two rubber gasket strips and clamping pressure. A United Sensor Corp PAD-6-KL pitot-static probe (11) outputs differential pressure to pressure Transducer A (12) to measure fluid velocity, and two pressure ports (7,8) output total pressure to pressure Transducer B (13) upstream and downstream of the fin array to measure pressure drop.

Both are Omega PX-277 differential pressure transducers, outputting analog data through a NI USB-6002 data acquisition module (14) to a Labview program running real-time data collection and monitoring (15).2318Computational TestingSimulations were run in ANSYS Workbench 15.0.7 using the FLUENT to handle three- dimensional CFD calculations. Fins were defined as aluminum with a constant temperature input at the bottom surface of the array, using array geometry and fluid flow characteristics equal to the experimental trials. A comparison of the experimental and simulated pressure drop results serves as validation for the accuracy of the simulated fluid flow, as well as a confidence measure for the simulated heat transfer results.Meshes were created using the ANSYS Meshing module. To minimize the necessary computation resources and time, models were sliced in half bilaterally and symmetry conditions used along the sliced surface.

To minimize the necessary computation resources and time, models were sliced in half bilaterally and symmetry conditions used along the sliced surface. Automatically generated unstructured grids were used to satisfy the complexity of closed-channel fluid flow through an array of obstructions. The meshes were refined around critical areas of low cross-section, such as between pins, and at areas of high aspect ratio, such as at the angled tip of some fins. Inflation layers were utilized to better capture the convective heat transfer effects of turbulent fluid flow near the pin surface, as well as to increase the resolution of internal conductive heat transfer in the pin itself (Figure 5). Initial experiments were performed to determine the necessary mesh resolution for grid-independence, which was defined as a less than 2% difference in overall heat transfer rate between trials.

Initial experiments were performed to determine the necessary mesh resolution for grid-independence, which was defined as a less than 2% difference in overall heat transfer rate between trials. Meshes used in this work utilize 300,000 – 1,000,000 computational nodes depending on array complexity.Figure 5: Top-view of meshing detail, illustrating the variable unstructured grid design and inflation layer application. (a) straight 0.6AR fin, (b) low-taper SEF, (c) high-taper 0.6AR.Solutions were calculated using two linked models. The energy equation was solved for conductive heat transfer within the solid fin array, and convective heat transfer between the fins and the fluid. The standard k-ε turbulence model was applied to calculate fluid flow characteristics, using second-order upwind spatial discretization for pressure, momentum, turbulent kinetic energy, turbulent dissipation rate, and energy parameters.

While the first-order algorithms result in faster convergence, second-order is more accurate in cases where flow is not aligned with the grid [22, p. 1409], as it is here with an unstructured grid and obstructed flow.23192The SIMPLE pressure-velocity coupling scheme was used, and the gradient discretization was least squares cell based.Experimental AnalysisExperimental testing produced 104 pairs of differential pressure data points per trial, ten trials per fin array, for 21 different fin array geometries. The following process was applied to each data set.First, the set of analog voltage data points Vin were converted to pressure Pin using the respective full scale pressure of each transducer PFS.

The following process was applied to each data set.First, the set of analog voltage data points Vin were converted to pressure Pin using the respective full scale pressure of each transducer PFS. Assuming steady incompressible flow, the resulting pressure from the pitot-static probe is converted to velocity vp-s using air density ρ at environmental temperature and pressure [23].𝑃𝑖𝑛 =𝑉𝑖𝑛 10𝑃𝐹𝑆(1)𝑣𝑝−𝑠 =2𝑃𝑖𝑛,𝑝−𝑠 𝜌(2)The standard deviation σ of each data set was calculated, and then an average taken to condense the 10 seconds worth of data into a single value. The error for each data point was calculated using the root-mean-square of the standard deviation and listed 1% full-scale error of the transducers.𝐸𝑟𝑟𝑜𝑟 = √𝜎2 + 0.01𝑃𝐹𝑆(3)The hydraulic diameter of the inlet is calculated using the height, a, and width, b, of the test section aluminum tubing.

This was used along with inlet velocity and kinematic viscosity ν to calculate the Reynold’s number Re of the flow entering the test section.𝐷ℎ,𝑖𝑛 =2𝑎𝑏(𝑎+𝑏)(4)𝑅𝑒 =𝐷ℎ,𝑖𝑛𝑣𝑝−𝑠 𝜈(5)Simulation AnalysisEach fin array was simulated individually with inlet fluid velocities of 1, 5, 9.5, and 14 m/s. Each simulation was configured to output a text file with total convective heat transfer rate Q, inlet fluid velocity vin, inlet pressure Pin, outlet pressure Pout, volume-weighted average temperature of the fluid volume Tfluid, and the area-weighted average wall temperature of the heat transfer surface Twall.First, the hydraulic diameter of the test section Dh,test was found using the internal fluid volume Vfluid and wetted surface area Afluid.

Pressure drop over the array ∆P was found using the difference between inlet pressure Pin and outlet pressure Pout.2320𝐷ℎ,𝑡𝑒𝑠𝑡 =4𝑉𝑓𝑙𝑢𝑖𝑑 𝐴𝑓𝑙𝑢𝑖𝑑𝛥𝑃 = 𝑃𝑖𝑛 − 𝑃𝑜𝑢𝑡Heat flux q was calculated from the total convective heat transfer rate and transfer area Awall, and was used to find the average convective heat transfer coefficient havg. This required the average pinwall temperature and average fluid volume temperature.𝑞 =𝑄 𝐴𝑤𝑎𝑙𝑙ℎ𝑎𝑣𝑔 =𝑞 𝑇𝑤𝑎𝑙𝑙−𝑇𝑓𝑙𝑢𝑖𝑑The average Nusselt number Nu was then found using the thermal conductivity k. This is a ratio of convective to conductive heat transfer across the solid-fluid boundary, and an indirect measure of average turbulence in the flow field. Higher values are desired for effective heat exchanger performance. The total pressure loss or friction coefficient f was calculated using the constant air density ρ and inlet velocity.

Higher values are desired for effective heat exchanger performance. The total pressure loss or friction coefficient f was calculated using the constant air density ρ and inlet velocity. This is a measure of relative pressure loss with respect to inlet flow conditions.𝑁𝑢𝑎𝑣𝑔 =ℎ𝑎𝑣𝑔𝐷ℎ,𝑡𝑒𝑠𝑡 𝑘𝑓 =2𝛥𝑃 2 𝜌𝑣𝑖𝑛Finally, the specific friction loss εf is calculated as a ratio of the average friction coefficient and average Nusselt number [11], [13]. This is an overall performance index, which indicates the pressure loss required to achieve an equal amount of heat transfer capability through the fin array. Minimizing this parameter equates to a more efficient and effective heat exchanger.𝜖𝑓 =𝑓 𝑁𝑢𝑎𝑣𝑔ResultsComparison of Experimental and Simulated Pressure Drop ResultsFigures 6-7 display a comparison of pressure drop between experimental results and simulated results for each fin array, grouped by geometry family.

It is observed that the simulated results increasingly diverge from the experimental results as Reynold’s number2321(6)(7)(8)(9)(9)(10)(11)increases. This is a result of the simulations, by the design and density of the mesh, and the second-order upwind governing k-ε equations, not capturing the full chaotic behavior of fluid flow in highly turbulent regions within the array [24].In all cases the experimental pressure drop was greater than the simulated values. One of the major contributors to this disparity was likely the simulated effects of fluid heating on viscosity and turbulence generation. From Re = 0 – 20,000 the data sets are highly similar, with mismatch error between 0–12%. In the more turbulent regime between Re = 20,000 – 50,000, higher mismatch errors are observed over a broader range, between 5–30%.

In the more turbulent regime between Re = 20,000 – 50,000, higher mismatch errors are observed over a broader range, between 5–30%. While less significant error may be required for design purposes, the quantitative correlation and relative similarity between experimental and simulated data sets in this work provide a strong basis for confidence in the validity of the simulated fluid flow characteristics.Figure 6: Comparison of experimental and simulated pressure drop across the 0.8AR geometry family.2322Figure 7: Comparison of experimental and simulated pressure drop across the swept back geometry family.Simulation Results Heat TransferFigure 8 shows the Nusselt number data sets for each geometry family as well as the circular test case. This metric not a direct measure of heat transfer, but of heat transfer enhancement due to the fins.

This metric not a direct measure of heat transfer, but of heat transfer enhancement due to the fins. A best fit line compiled by Ciha from a survey of ten studies on similar fin geometries [25] is presented along with the results of this work, on a log-log scale for direct comparison.The density of information makes it difficult to glean useful associations between test cases from this plot, but it does serve as an effective comparison between this work and previous works in the literature.

19 of the 21 fin arrays tested align well with this survey data, while the 0.6AR straight and 0.8AR straight arrays are outliers.Contrasting these two with the circular fin array, which presents less effective heat transfer than the best fit line, the cause of this disparity was a combination of three effects: the high level of flow obstruction in the straight configuration, an increase in length travelled along the fin before flow separation commensurate with the length of the tail, and more efficient heat transport from the base to the fin tip related to cross-sectional radially symmetry. The 0.6AR straight array creates more efficient flow separation but less efficient conductive heat transport, while the 0.8AR straight array creates the opposite effects.2323Figure 8: Comparison of average Nusselt number data for all fin arrays.

The best fit line is reported from a survey similar pin fin arrays and flow regimes [25].Figure 9: Overall comparison of array heat transfer capacity using the sum of heat transfer rates at four Re inputs, from best to worst. ✝ Values calculated from the best fit equation rather than directly from simulation data.2324Figure 9 shows average heat transfer rate for each fin array, area-weighted over the full heat transfer surface in each simulation case, combining sets of data at four inlet flow velocities. These data are presented without normalization to any geometry or fluid parameter, and represent the total amount of energy flowing from the pins to the fluid through convective heat transfer. While the Nusselt number is an effective metric for overall efficiency of heat exchanger arrays, Q represents a pure energy transfer basis for comparison. If pressure drop is not a limiting factor in design, this represents the performance scale for the tested fin arrays.

If pressure drop is not a limiting factor in design, this represents the performance scale for the tested fin arrays. Moving from best to worst constitutes moving through the straight arrays, to the low taper arrays, and finally to the high taper arrays.Heat transfer results follow the theoretical geometry correlations in these data expected from previous work [6], [11], [13], [18], [26]. Higher flow obstruction induces more turbulent flow and accelerates the fluid through narrower inter-fin channels, and longer fins tails delay flow separation. These phenomena are illustrated in Figures 10-11, which show triptych contour plots of the 0.2AR, 0.6AR, and 0.8AR fin arrays for velocity and turbulent kinetic energy, respectively, on normalized scales.Figure 10: Normalized velocity contour plots of the fluid volume, taken on the horizontal midplane of each fin array.

✝ Values calculated from the best fit equation rather than directly from simulation data.Overall PerformanceIn the previous two sections, some physical mechanisms responsible for the relative differences in heat transfer and pressure drop among the tested fin arrays were discussed. Since these mechanisms affect both parameters in opposite ways, balancing the two becomes a design optimization problem based on fin geometry and array configuration. To quantitatively contrast fin arrays, the non-dimensional specific friction loss metric was introduced in [9]. This is the ratio of friction coefficient to Nusselt number, basically a measure of the pressure loss required to achieve a certain amount of heat transfer capability across the fin array, where lower values are desired. Figure 13 displays an overall comparison of the specific friction loss for each fin array, combining sets of data at four inlet flow velocities. This provides a holistic basis for inter- array comparison.

This provides a holistic basis for inter- array comparison. It can be seen that the 0.2AR high taper array is the most effective of those tested in the present work, and furthermore that every design was more effective than the standard circular fin geometry.2327Figure 13: Comparison of array performance using the sum of specific friction loss at four Re inputs, from best to worst. ✝ Values calculated from the best fit equation rather than directly from simulation data.Discussion Effect of Tapering on PerformanceIt is clear from the results of this work that, within each geometry family, increased taper angle correlates with more effective performance.

Figure 14 displays this trend in the SEF geometry family, and the causal physical phenomena as described in previous sections are visible in Figures 10-11.2328Figure 14: Overall performance of the SEF geometry family, characterized by specific friction loss.Effect of Varying Fin Dimensions Along the Streamwise AxisThis work represents a proof of concept for diameter changing fins along the length of a heat exchanger array. By varying the cross-sectional geometry of each fin as a function of its location in the fluid stream, both heat transfer efficiency and pressure drop are manipulated, and it was theorized that a favorable ratio of change in these parameters could increase overall performance of the array.

This theory has been preliminarily confirmed in this work, as results for the diameter change low taper array place it as the third best performing of the 21 tested (Figure 13).Figure 15 shows normalized turbulence contour plots for the straight and low taper diameter change arrays. It is observed that turbulence increases as it enters the fin array, reaches a maximum around the fifth row, and then steadily decreases. To increase the effectiveness of these geometries, turbulence should be generated in the initial rows and then maintained, and these results indicate the use of an over-aggressive diameter reduction scheme.2329Figure 15: Normalized turbulence contour plots of the fluid volume, taken on the horizontal midplane of each fin array. a) Diameter change straight, b) Diameter change low taper.Comparison of Linear and Nonlinear TapersThe comparisons so far made all been used linear tapers, with a constant draft angle between the base and tip of the fin.

Considering that the increased performance observed in tapered fins is derived from a greater match between the vertical cross-sectional profile and temperature gradient, and that this gradient is likely nonlinear due to geometry and boundary layer effects at the base, a nonlinear taper has been theorized to effect a higher overall performance in this type of pin fin array [27]. Two nonlinear profiles were tested in this work as a proof of concept for future investigation, both concave and convex. Figure 16 shows the performance of these two arrays compared with the linear tapered version, on a log-log scale with best fit lines shown. It is observed that the convex geometry performance is at best equal to the linear version, with a significant decrease at lower Reynold’s number flow.

It is observed that the convex geometry performance is at best equal to the linear version, with a significant decrease at lower Reynold’s number flow. The concave geometry performance is significantly greater than the linear version at all tested flow regimes.2330Figure 16: Performance comparison between the 0.6AR array geometry with linear high taper, concave taper, and convex taper, including calculated best fit lines.Comparison of Double and Triple Repeating Fin PatternsFigure 13 shows that pin fin arrays with staggered rows repeating every three instances are viable candidate designs for further study on an overall performance basis. These arrays were modelled on the 0.6AR low and high taper arrays, with equivalent streamwise and spanwise fin spacing, but rearranged to form the triple row pattern. This led to a decrease in the total number of pins over the array and subsequently more volume in the inter-fin region, so direct quantitative comparisons require further study.

This led to a decrease in the total number of pins over the array and subsequently more volume in the inter-fin region, so direct quantitative comparisons require further study. Still, normalized velocity contour plots for the double- and triple-repeating arrays of 0.6AR pin fins are shown in Figure 17, and allow us to quantitatively compare the fluid flow in each configuration.It is observed that the triple row pattern produces a less linear configuration of fluid flow, with a higher degree of spanwise motion. Impinging flow on the front of each fin is shed to the inter-fin region where, instead of continuing straight through to the outlet, encounters the shoulder of another fin. This cascade effect served to increase both turbulence and heat transfer.2331Figure 17: Normalized velocity contour plots of the fluid volume, taken on the horizontal midplane of each fin array.

This cascade effect served to increase both turbulence and heat transfer.2331Figure 17: Normalized velocity contour plots of the fluid volume, taken on the horizontal midplane of each fin array. a) 0.6AR low taper, b) (0.6AR) triple row low taper.Achieving Optimal PerformanceThe overall highest performing geometry observed in this work is the 0.2AR high taper array (Figure 13). The high aspect ratio teardrop shape effectively reduces turbulence and vortex shedding in the wake of the fin, delays boundary layer separation, and the high taper increases heat transfer performance at the tip of the fin while allowing greater fluid flow.

As a result, the overall pressure loss experienced by this fin array was minimized, with respect to the proportional increase in heat transfer efficiency.The general trends observed by previous work are that increasing aspect ratio of teardrop fins increases their overall performance [11], [13], and increased array density can lead to increased overall performance [5], [9]. In this work it was observed that overall performance can be increased by using high taper angles, nonlinear taper profiles, and variations on the staggered array pattern, as well as reducing fin diameter along the streamwise axis.Following these trends to their conclusions, using 0.1AR or similar extremely high aspect ratio fins, utilizing a high taper angle with nonlinear concave profile, in a more condensed version of the triple or quadruple row repeating layout pattern, with variable diameter along the length of the array, could achieve more optimal results than those found in this work.

Mechanical design considerations would need to be addressed; the fin tail ends may be fragile to debris or highly turbulent flow, and the arrays may become impractically long to maintain proper fin spacing.2332Future WorkDiameter Changing Fin ArraysFuture work in this area should include alternative diameter reduction schemes should be investigated, both less and more aggressive than 5% per row. It should also include varying the aspect ratio and taper parameters of fins along the streamwise length, instead of solely the fin diameter. In theory it should be possible to discretize this type of array into quantitative parameters, which could be inserted into a machine learning algorithm to intelligently optimize the shapes, locations, and characteristics of each individual pin for a specific flow regime. The outputs of this algorithm would be used as design parameters for computational testing, similar to the process employed by Wu et al.

The outputs of this algorithm would be used as design parameters for computational testing, similar to the process employed by Wu et al. utilizing genetic algorithms [28].Nonlinear and Variable Cross-Section Tapered FinsFuture work in optimization should be conducted, defining fin profile based on a theoretical and simulated analysis of steady-state vertical temperature profiles along the length of each fin. Malekzadeh et al. introduced an accurate two-dimensional numerical tool for analysis of nonlinear fins in [27], which could be employed or extended to 3D for use in the design stage. In addition to the taper profile, varying the cross-sectional geometry as a function of fin height should also be investigated.

In addition to the taper profile, varying the cross-sectional geometry as a function of fin height should also be investigated. Optimization using both methods could be further accomplished through iterative machine learning or design of experiment testing.Triple Row Repeating Fin ArraysThe increase in fluid-fin interaction caused by the triply staggered rows is responsible for enhanced heat transfer within the fin array. Future work should be conducted on these triple row repeating arrays, investigating the effects of streamwise and spanwise spacing on tapered pin fins of various cross-sectional geometries.Teardrop Fin Arrays at High Reynold’s NumbersFuture work should be conducted to confirm the relative performance of straight teardrop pin fins at high Reynold’s number when compared to tapered analogues.

The lack of such a relationship in the SEF geometry family leads this author to believe it may have been an artifact derives of some part of the simulation design.SummaryTo study the effects of novel tapered pin fins shapes and arrangements on heat transfer efficiency, pressure loss, and overall performance, 21 fin arrays were designed and implemented across eight geometry families with varying fin cross-section, taper angle, taper profile, and array patterns.Experimental testing was carried out on a prototype open-loop wind tunnel system using differential pressure sensors to measure pressure drop across the test section directly, and inlet flow velocity indirectly using a pitot-static tube. The fin arrays were fabricated in Nylon 12 using laser sintering, and data gathered using at 0 < Re < 50,000.2333Computational fluid dynamics simulations were designed in ANSYS to model pressure drop and heat transfer across heated aluminum versions of the same fin arrays.

Experimental pressure drop data was used as validation for the simulated fluid flow, and as a confidence metric for the accuracy of the simulated heat transfer.The experimental and computational data were reduced and analyzed using Matlab, defining non-dimensional metrics for holistic comparison between the disparate array geometries and layouts. These metrics were used to compare heat transfer efficiency, pressure loss characteristics, and overall balanced performance between fin arrays.Comparisons were made between members of each geometry family, as well as between geometry families, to illuminate effective means for increasing overall heat exchanger performance.

Topics for future work and potential methods of investigation were suggested.Conclusions1.Teardrop shaped pin fins effectively increase the performance of pin fin arrays by delaying boundary layer separation, and wake field turbulence.2.High aspect ratio teardrop fins are desired to maximize these effects and attain the greatest overall performance.3.Tapered pin fins in any cross-sectional geometry are more effective than straight pin fins in equal flow regimes.4.Of the 21 tested fin arrays, the 0.2AR high taper array was the overall highest performing geometry.5.Varying fin dimensions along the streamwise axis can effectively increase the performance of drop-shaped pin fin arrays. These configurations should be investigated in more detail.6.Using nonlinear concave taper profiles can effectively increase the performance of drop- shaped pin fins.

These configurations should be investigated in more detail.6.Using nonlinear concave taper profiles can effectively increase the performance of drop- shaped pin fins. This type of fin geometry should be investigated in more detail.7.Alternative fin placement configurations, such as triple row repeating staggered arrays, can effect more spanwise fluid movement and ultimately increase the performance of drop-shaped pin fin arrays. These configurations should be investigated in more detail.References[1] C. L. Chapman, S. Lee, and B. L. Schmidt, “Thermal Performance of an Elliptical Pin Fin Heat Sink,” Proc. 1994 IEEE/CHMT 10th Semicond. Therm. Meas. Manag. Symp., 1994, 1994.[2] O. Uzol and C. Camci, “Elliptical Pin Fins As an Alternative To Circular Pin Fins for Gas Turbine Blade Cooling Applications, Part 1,” ASME Turbo Expo 2001 Power Land, Sea, Air, vol. 3, 2001, 2001.[3] G. J. VanFossen, “Heat-Transfer Coefficients for Staggered Arrays of Short Pin Fins,” J.Eng.

VELOPMENT OF POWDER BED FUSION ADDITIVE MANUFACTURING TEST BED FOR ENHANCED REAL-TIME PROCESS CONTROLM. L. Vlasea*, B. Lane*, F. Lopez*, S. Mekhontsev*, A. Donmez* * National Institute of Standards and Technology, Gaithersburg, MD 208991AbstractLaser powder bed fusion (PBF) is emerging as the most popular additive manufacturing (AM) method for producing metallic components based on the flexibility in accommodating a wide range of materials with resulting mechanical properties similar to bulk machined counterparts, as well as based on in-class fabrication speed. Although this approach is advantageous, the current limitations in achieving predictable and repeatable material and structural properties, geometric and surface roughness characteristics, and the occurrence of deformations due to residual stresses results in significant variations in part quality and reliability. Therefore, a better understanding and control of PBF AM processes is needed.

Therefore, a better understanding and control of PBF AM processes is needed. The National Institute of Standards and Technology (NIST) is developing a testbed to assess in-process and process-intermittent metrology methods and real-time process control algorithms, and to establish foundations for traceable radiance- based temperature measurements that support high-fidelity process modeling efforts. This paper will discuss functional requirements and design solutions to meet these distinct objectives.1IntroductionLaser powder bed fusion (PBF) additive manufacturing (AM) processes, are highly versatile manufacturing strategies, demonstrating a range of competitive advantages over conventional manufacturing in terms of enabling the production of customized parts with intricate internal lattice features and complex external geometries [1,2] without the use of dedicated tooling [1], essentially reducing the time between digital design and fabrication [3].

In addition to the customization of geometrical features, laser PBF processes have the potential of producing lightweight or weight-optimized lattice structures with suitable mechanical properties [4] and multi-material composition [2], with minimal material loss. Although the parts produced using laser-based AM approaches tend to have embedded pores, their densities can reach almost 100%, having mechanical properties equaling and at times exceeding those of conventionally manufactured counterparts [1,5,6]. In the new generation of digital design considerations, a breakeven point exists where the complexity, weight optimization, and low production demands of parts favor AM approaches [6], with industry contenders ranging from the medical field, to aerospace applications.

The benefits of laser-based PBF have sparked interest among the research community and industry, with an increased focus on strategies to overcome some of the existing challenges related to part quality, consistency, and repeatability [7].1 Official contribution of the National Institute of Standards and Technology (NIST); not subject to copyright in the United States. The full descriptions of the procedures used in this paper may require the identification of certain commercial products. The inclusion of such information should in no way be construed as indicating that such products are endorsed by NIST or are recommended by NIST or that they are necessarily the best materials, instruments, software or suppliers for described purposes.According to literature in the field, PBF processes have a high degree of sensitivity to disturbances and input process parameters such as laser power, laser scanning path, and scan speed.

The vast number of controllable and predefined process parameters as summarized by Mani et al. [7] have a complex influence over the transient thermal behavior of the melt pool, often resulting in unexpected defects (pores) [8], high surface roughness [3,7], thermal cracking and delamination [9], unintended anisotropic mechanical and physical properties [1,3,10], as well as anisotropic shrinkage [11]. Pores can be traced back to the high temperature input associated with laser AM that may cause material vaporization and spatter during the dynamic laser-substrate interaction [3]. Undesirable porous regions can also be associated with the so- called balling phenomenon, which can occur due to poor wettability of the previous layer resulting in molten material not binding to the underlying substrate. This type of process defect has been linked to insufficient local laser power or high scan speeds [10].

This type of process defect has been linked to insufficient local laser power or high scan speeds [10]. Surface roughness and surface defects can become severe, especially in so-called overhang zones, where sections of the in-process layer are supported by powder. In these areas, the laser-powder interaction causes local overheating due to poor heat conduction and dissipation, resulting in an over-sized melt pool geometry [12]. Furthermore, high cooling rates and repeated thermal loading caused by latent heat of fusion, adjacent laser scanning paths within the layer, as well as heat imparted from subsequent layers, have a significant effect on local transient temperature distributions and temperature gradients, triggering rapid heating, melting, and solidification cycles [2]. These effects cause high residual stresses in the part [13,14] resulting in curling of surfaces, microstructure heterogeneity [15], and dimensional deviations of the final part [11,16].

These effects cause high residual stresses in the part [13,14] resulting in curling of surfaces, microstructure heterogeneity [15], and dimensional deviations of the final part [11,16]. It is important to reduce thermal stresses, as they cause premature part failure and defects [12].The vast number of process parameters interacting to create a dynamic melt pool environment increase the complexity level of laser PBF manufacturing approaches. Current strategies for determining optimal process parameters of emerging materials or refinements for existing materials still rely on expensive and time-consuming trial and error iterations [7,9,17]. To minimize the need for such iterations, researchers are working on developing high fidelity physics-based models of AM processes to narrow down the input parameter range.

To minimize the need for such iterations, researchers are working on developing high fidelity physics-based models of AM processes to narrow down the input parameter range. However, these models rely on accurate input material properties such as diffusivity, reflectance, and emittance of materials in multiple states of aggregation (powder, liquid, solid), which are currently unavailable [18,19]. These models also require traceable process measurements for fine-tuning and validation. In commercially available machines, once appropriate process parameters have been identified, they are typically kept constant, often leading to part defects and process instability as the process becomes more sensitive to disturbances [1,20]. In order for laser-based AM technologies to become an integral part of mainstream manufacturing, the produced parts have to adhere to stringent quality standards, with repeatable and reliable characteristics [4,5].

To achieve this goal, it is important to advance the science behind developing appropriate metrology strategies to aid in refining our understanding of the relationship and relative sensitivities between input process parameters, process signatures, and product qualities [7], and integrate the knowledge gained into effective quality monitoring and control strategies [5].2 Measurement organizational structure for controls implementationWhile there have been multiple research efforts and even commercial implementation in sensing, monitoring, and controlling metal-based AM processes, there is still significant research required before they can become commonplace on commercial AM systems.

In the literature, most papers either investigated the correlations between process input parameters and product qualities [20,21] or exclusively focused on defining relationships between process parameters and melt pool characteristics as process signatures [10,12,13,18], with some works concentrating on in- situ process control strategies [1,3,5,15,17,19,22]. In their review work, Mani et al. [7] reviewed metrology approaches, as well identified the correlations between process parameters, process signatures, and product qualities with a direct relevance to developing strategies for in-situ monitoring and control. To further establish the foundations for AM process control, the process parameters, process signatures, and product quality were sub-categorized according to the abilities to be measured and/or influenced as control parameters in the process as shown in Figure 1 [7].

Prior to this work, process parameters and signatures relations from literature had not yet been logically related to product quality in a broader, general fashion.Figure 1 The hierarchy of process parameters, process signatures, and product qualities as identified by Mani et al. [7]This current work focuses on providing the relationship map between measurand, measurement method, and control strategies in laser PBF through references to current research efforts, as wellas outlining limitations and potential opportunities in advancing intelligent controls in AM. An organizational structure for control strategies, presented in context with the previously established hierarchy of parameters [7] is proposed as seen in Figure 2. In this figure, the measurand is defined as the parameter information acquired through appropriate measurement technologies and processed into a suitable data stream for post-process analysis, in-situ process monitoring, and feedback process control.

As part of control strategies, four types of approaches have been identified and will be discussed in further detail: pre-processing for predictive control, in-situ defect or fault detection and handling, in-situ continuous feedback control, and signature- derived control through plant models or simulations. Another category of feedback-dependent strategy closely correlated with process control outcomes is part qualification through in-process monitoring and analysis, which is an emerging field in AM part certification and quality control [18].

This type of technique takes advantage of the layer-by-layer manufacturing approach to perform continuous process-signature analysis and layer-intermittent inspection of layer qualities.Figure 2 Measurement organizational structure in the context of possible controls strategies for PBF AM2.1 Pre-processing for predictive controlPresently, there is an active interest in refining AM processes for fabricating products with heterogeneous or functionally graded structural, compositional, and mechanical properties with applications in industrial, aerospace, and biomedical fields [23], as well as a continued interest in refining AM processes for producing products with consistent and repeatable part qualities.

In this context, pre-processing for predictive control can have a multi-layered architecture that first considers (a) digital pre-processing for three-dimensional topology design, and subsequently considers (b) an input parameter optimization stage to achieve the desired design configurations.The control strategy for digital pre-processing of three dimensional parts for compositional and topological optimization, focuses on the digital design and characterization using modelingapproaches. Modelling should consider discrete and stacked lattice structures composed of multi- material and/or multi-structural cellular features to investigate configurations that can meet design criteria as well as compute a range of initial input process parameters for the AM process.

Previous modeling work has typically focused on balancing the tradeoff between structural topology for lightweight design and resulting mechanical properties [24,25], as well as modeling the relationship between process parameters and resulting thermal behavior to infer refinements of the input parameter space based on trends to avoid high thermal loading and associated potential defects [2,11,14,26–28].

To the authors’ knowledge, no work has been done on merging structural topology optimization and thermal modeling into a coherent digital pre- processing algorithm at the 3D scale.Secondly, there is an opportunity to investigate optimization strategies that can enable further refinement of input process parameters by taking into account experimentally measured process parameters, process signatures, and product quality data from previously manufactured batches of similar products to continuously improve geometric shape fidelity and account for drifts in machine performance and raw material batch variation.

For example, one strategy could be to continuously monitor the relationship between the designed 3D digital model, the 3D in-line reconstructed model (from thermal, acoustic, layer imaging, laser profilometry, contact image sensing, infrared detection data), and the 3D generated scan of the final part to generate pre- processing predictive control strategies for better part quality, process speed, and reliability. One example is the work done by Ning et al.

One example is the work done by Ning et al. [11], where an experimentally-derived model was developed to correlate 2D in-layer geometric shape to shrinkage, and furthermore, the model was used in a feedback scan compensation algorithm to pre-compute laser scan-speed compensation factors resulting in increased dimensional accuracy of subsequent parts.2.2In-situ defect or fault detection and handlingIn these schemes, the controller responds to defects or faults by triggering discrete events that interrupt normal control routines and activate a corrective action to avoid process instability which may lead to catastrophic failure of the manufactured part. While some errors or faults can be compensated via continuous feedback control, some errors require separate error handling routines, which are the focus of this section. Overlaps between the two control strategies should exist, as they are complementary and share the same measurement parameter space.

Overlaps between the two control strategies should exist, as they are complementary and share the same measurement parameter space. In this context, process errors can be related to non-uniform powder layer spread due to damage of the re-coating mechanism [5], collisions between the re-coater and super-elevations [13], powder exhaustion [5,18], or contaminants in the powder substrate. Process errors can also be related to defects in the track or layer such as unintended porosities and high surface roughness due to un- melted particles, material ejection during melting, and balling effects [3,9,29], as well as excessive shrinkage and curling due to residual stress [11,13,14].

Other errors relate to machine parameters such as laser beam delivery and chamber environment [19], and can be monitored and controlled such that the process is interrupted until normal operation can resume.Research focusing on the quality of powder spread has generally made use of imaging systems such as a monochrome charge-coupled device (CCD) camera and dark field illumination with an angled viewpoint to highlight surface powder disturbances for qualitative measurements [20], aswell as for quantitative evaluation of powder profile [5] to evaluate re-coater blade extent and location of damage. One possible controller corrective action is identification of areas of inconsistent powder spreading before the deposition process begins and re-positioning the part at a more convenient location in the build space. If the powder spread error occurs during the build phase, attempting to re-spread a new layer or dynamically adjusting or replacing the re-coater blade could also be feasible.

If the powder spread error occurs during the build phase, attempting to re-spread a new layer or dynamically adjusting or replacing the re-coater blade could also be feasible. When detecting process errors related to pores or part surface roughness, research has mainly focused on thermal process signature measurements and the correlation between abnormal measurements and defect outcome using either photodiode [5] or pyrometer [10] measurements, or microbolometer detectors [18]. For instance, Krauss et al. [18] have focused on detecting porous defects using thermal monitoring via an uncooled microbolometer detector by considering both spatial and temporal signatures of the heat affected zone to detect pores up to 100 μm in size. Craeghs et al.

Craeghs et al. [5] have coupled a photodiode signal with melt pool area measurements obtained from using a complementary metal-oxide semiconductor (CMOS) camera to detect abnormal trends which can be correlated to porosities, surface roughness errors, layer thickness inconsistency errors, and gradual part quality deterioration. Most relevant research in this area focuses on defect, error, or fault detection, evaluating trends for offline refinement of input parameters, with little work done on in-line error detection coupled with corrective actions.2.3In-situ continuous feedback controlFeedback process control schemes to-date allow for controllable process input parameters to be continuously modulated based on feedback from measurable process signatures or product qualities via decisions from a controller algorithm [19].

Additional control loops should be implemented to monitor the input process parameters to ensure that the equipment is actually running according to specified command input signals.The first implementations of feedback control in metal-based AM originated in the laser surface alloying and laser cladding industries. Römer et al. [30] proposed a temperature and melt pool controller for the laser alloying process fed with thermographic measurements. Similar approaches were followed for the laser cladding process by controlling clad height [31,32], melt pool temperature [15,33], or melt pool width [34]. Feedback control has also been used in direct energy deposition (DED) processes, to control melt pool temperature [35,36] or melt pool area [37] and similarly in SLM, melt pool parameters [1,5,17,22] and plasma plume measurements [3]. All of these approaches used thermal measurements obtained with CCD, CMOS, or infrared (IR) cameras.

All of these approaches used thermal measurements obtained with CCD, CMOS, or infrared (IR) cameras. There is a consensus in the need for thermal measurements for feedback control and the utilization of proportional-integral-derivative (PID) controllers is a popular approach [15,21,22,30–35,37], as well as other emerging implementation strategies such as model predictive control [36]. The design of said controllers is often based on heuristics [15,35,37,38], or on plant models obtained through system identification [12,21,30–33,36], and rarely on physics-based models [34].

The design of said controllers is often based on heuristics [15,35,37,38], or on plant models obtained through system identification [12,21,30–33,36], and rarely on physics-based models [34]. The development of such high fidelity physics-based models of AM processes rely on accurate measurements of input material properties such as diffusivity, reflectance, and emittance of materials in multiple states of aggregation (powder, liquid, solid), measurements which are difficult to obtain, and are the incentive behind the development of a new AM metrology test bed at NIST.Despite all the work in continuous feedback control, there is still room for improvement, especially in thermal metrology and controller design. Previous implementations of thermal imaging often oversimplified calibration, ignored camera saturation, or assumed constant surface emissivity. All of these factors may result in erroneous measurements that insert error in the feedback control loop.

All of these factors may result in erroneous measurements that insert error in the feedback control loop. Also, system identification approaches require extensive testing to identify the dominant dynamics of the plant. Results are often not extendable to other materials, or geometries, limiting their applicability.2.4 Signature-derived control through plant models or simulationsSignature-derived control schemes rely on results (estimations) derived from process modeling or simulation algorithms that use in-situ feedback signals as inputs to compute or infer immeasurable quantities. These derived quantities (or estimations) can then be used as further inputs to a control algorithm. This implies that simulations may run in parallel with the controller or that process maps and lookup tables are dynamically accessible. In general, estimation is often ignored in feedback control for metal-based AM.

In general, estimation is often ignored in feedback control for metal-based AM. This may be attributed to two factors: 1) the lack of appropriate models that map measurable process parameters to immeasurable process signatures, and 2) the fast dynamics of the processes that force all measurements to be gathered and processed at very high sampling frequencies. In the literature, the preferred trend is to find a way to measure the quantity of interest directly rather than to use available measurements to infer it [19]. As an example, the work done by Aggarangsi et al. [8], where a thermomechanical finite element analysis (FEA) model was developed to estimate residual stresses within a thin wall and process maps were devised to correlate stress, melt pool depth, and laser power, could be further enhanced as signature-derived control. Another example is the model-based controller developed by Devesse et al.

Another example is the model-based controller developed by Devesse et al. [34], where temperature measurements were used as feedback to estimate melt pool size and to ultimately control the laser power. Other developed models based on experimentally derived parameters [9,39] show promise in this area. Although often ignored, there is a wealth of opportunity in developing estimation methods or models for feedback control.2.5 Part qualification through process monitoring and analysisThese control schemes enable automated qualification of the final part after the build is complete. The controller decision is the culmination of multiple measurements, or a single measurement at the end of the build, and quantifies product quality form in terms of geometry, material structure, surface finish, or porosity characteristics. There are currently no recognized industrial or research solutions that address this need. In their report, Mani et al.

There are currently no recognized industrial or research solutions that address this need. In their report, Mani et al. [7] have identified a comprehensive list of in-process layer geometry measurements [13,40] and post- processing measurements of part quality applicable to dimensional accuracy [41], surface quality [42], mechanical properties [42,43], residual stress [44], porosity and density [45], and fatigue [46], that can help in part qualification. These product qualities have been studied by various research groups and correlated to input process parameters or with process signatures.

However the development of heuristic in-process or process-intermittent metrology approaches for part qualification has not been implemented and can prove to be an interesting research opportunity.3 Proposed additive manufacturing metrology test bedIn recognition of the possible opportunities in furthering research into monitoring, metrology, and process control for laser powder bed fusion, NIST is developing a SLS AM testbed with a dual purpose: (1) to assess in-process and process-intermittent metrology methods and real-time process control algorithms aimed to enhance process repeatability, reliability, and part quality, and (2) to establish foundations for traceable radiance-based temperature measurements that support high-fidelity process modeling and model validation efforts, as well as to study methods for real-time thermometry in the production environment.

It is anticipated that development of this testbed will contribute to the development of the next generation of AM mechatronic systems that would empower AM for widespread use for cost-effective functional part production. This section will introduce the functional requirements and some of the design solutions considered thus far to meet these two distinct objectives.3.1 General functional requirements and design solutionsThe system has been designed to operate in two modes, so-called Build Mode, used to meet the demands of purpose (1) as stated above, and Radiance-based Metrology Mode, to address purpose (2). In Build Mode, the system should be capable of basic part production functionality, where the re-coater arm is able to spread powder from a feed bed area to a build bed, and a laser system directs the beam to the appropriate location.

The build area has a temperature-controlled heated plate, with a temperature target of 80 ºC, with the capability of adding a high temperature build chamber-insert capable of heating the build plate and chamber walls to at least 250 ºC. The build plate can be manually adjusted to correct for tilt in the direction of powder spreading using a screw-activated pivoting kinematic clamp assembly. The re-coater arm allows for adjustment of the recoating blade tilt along the direction perpendicular to the direction of spread. Table 1 summarizes the design requirements for the build and feed compartments.

The Build Mode design should also accommodate for the appropriate sensors and instrumentation used to support in-process and process-intermittent metrology, as well as development of control strategies.Table 1 Design functional requirements for Build ModeComponentEffective SpecificationsFeed bed – height180 mmBuild bed – height (low temperature)90 mmBuild bed – height (high temperature)25-30 mmFeed bed – size150 x 150 mmBuild bed – size (low temperature)150 x 250 mmBuild bed – size (high temperature)Φ100 mmTemperature build bed (large chamber)80 ºC (175 ºF) bottomTemperature build bed (small chamber)250 ºC (482 ºF) (chamber)Build bed tilt alignment screw (x-axis)1.3 deg/revLaser power, continuous500 WattMaterialsmetal/alloy powdersIn Radiance-based Metrology Mode, the laser beam has to interact not only with the powder substrate in the build bed, but also with a range of calibration standards: an IR standard source, visible near IR (VIS-NIR) standard source, as well as with a VIS-NIR reflectance standard, with reflectometry measurements taken under an integrative reflectometer dome.

To achieve this task, the entire powder bed assembly was designed to be able to translate with high accuracy between various build and metrology positions. The powder assembly system, along with the actuation mechanisms is referred to as the powder bed carriage. The integrative dome is designed to be retractable by allowing reciprocating motion, moving in an out of the laser path on command.(a)(b)Figure 3 (a) Conceptual design shown in Build Mode, where the build area is aligned with the air blade, and the in-process and process-intermittent metrology methods and real-time process control algorithms can be tested. (b) Conceptual design shown in Radiance-based Metrology Mode, with the carriage moved such that the IR standard source is in position, under the metrology integrated dome.3.2 Proposed metrology instrumentation for monitoring and control opportunitiesThe system is currently transitioning from the design stage, to implementation.

The metrology instrumentation will include strategies for in-line coaxial continuous monitoring for thermal measurements using a single point high-speed filtered photodetector, as well as a coaxial IR imaging system to detect melt pool geometry. Vision-based systems will be integrated to monitor continuously or intermittently the powder spread quality, as well as estimating the layer product qualities such as geometry, pores, deformations due to residual stress, and general part photogrammetry. A modular re-coating arm will be integrated in the system to allow for testing of different sensing strategies including acoustic, as well as non-contact measurements such as laser profilometry, contact image sensing, and 2D laser displacement sensing. Process signature characterization will include strategies for radiometric measurements, emissivity measurements, and traceable thermal measurements, which will be used in improving and validating of modeling strategies.

Other sensing technologies are currently being explored.From a metrology standpoint, there is a recognized need to further the development of relationships between process parameters, process signatures, and part qualities, as well as exploring the sensitivities of those relationships through experiment and simulation [7], which will be one of the important goals of the newly developed AM test bed. Furthermore, there are opportunities in developing measurement strategies for a range of other less explored process signatures, such as laser ablation plume size, as well as spectral measurements in the process zone. An important consideration will be given to understanding the measurement uncertainty, as well as analysis of measurement error and traceability.

An important consideration will be given to understanding the measurement uncertainty, as well as analysis of measurement error and traceability. The first research phase using the newly developed test bed will focus on these metrology-related goals.The control strategy map illustrated in Figure 2 provides a large-scale outline of the possible controls approaches for powder bed fusion AM. Currently, control strategies rely mainly on correlations between melt pool geometry and temperature. The proposed schemes will focus on establishing a multi-control architecture, with the potential of encompassing pre-processing strategies, error detection and handling, continuous feedback control, and signature-derived control. These approaches will be explored in future work.4 ConclusionsThis work has focused on outlining the advantages and challenges of laser PBF AM, as well as understanding the need for developing a new generation of control strategies.

An organizational structure for possible controls strategies has been proposed, to encompass pre-processing predictive control, in-situ defect or fault detection and handling, in-situ continuous feedback control, and signature-derived control through plant models or simulations. In addition, part qualification through in-process monitoring and analysis has been discussed, which is an emerging field in AM part certification and quality control. The PBF AM testbed being developed at NIST aims to advance the current state of the art of metrology, monitoring, and control in the field. The system has been introduced in this work by describing the general functional requirements and design solutions, and the proposed vision for metrology and control.References[1] Craeghs T, Bechmann F, Berumen S, Kruth J-P. Feedback control of Layerwise LaserMelting using optical sensors. Phys Procedia 2010;5:505–14. doi:10.1016/j.phpro.2010.08.078.[2] Raghavan A, Wei HL, Palmer TA, DebRoy T.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDEVELOPMENT OF PRE-REPAIR MACHINING STRATEGIES FOR LASER-AIDED METALLIC COMPONENT REMANUFACTURINGXinchang Zhang*, Wenyuan Cui*, Leon Hill*, Wei Li*, Frank Liou*Department of Mechanical & Aerospace Engineering, Missouri University of Science and Technology, Rolla, Missouri, 65409, USAAbstractRemanufacturing worn metallic components can prolong the service life of parts that need frequent replacement but are extremely costly to manufacture, such as aircraft titanium components, casting dies. Additive manufacturing (AM) technology enables the repair of such valuable components by depositing filler materials at the worn area layer by layer to regenerate the missing geometry.

Additive manufacturing (AM) technology enables the repair of such valuable components by depositing filler materials at the worn area layer by layer to regenerate the missing geometry. In general, damaged parts would be inspected and pre-machined prior to material deposition to remove oil, residue, oxidized layers or defects located in inaccessible regions. Therefore, the motivation of this paper is to introduce pre-repair machining strategies for removing contaminated materials from damaged components and materials surrounding inaccessible defects to ensure that the target damage is repairable. The current research targets at common failures comprising surface indentations, erosion, corrosion, wear and cracking, and the machining strategies for each defect were proposed. Each strategy takes the 3D scanned damaged model as input and the cut-off volume around the defects is defined by using different approaches.

Each strategy takes the 3D scanned damaged model as input and the cut-off volume around the defects is defined by using different approaches. Pre-repair machining toolpath and program were generated based on the defined cut-off volume and finally, damaged parts were machined using the proposed strategies.Keywords: Pre-repair machining; Component repair; Defects; Additive manufacturing1. IntroductionAircraft components and casting dies are subjected to premature failure due to impact with foreign objects, thermal fatigue cracking, erosion, corrosion, wear, etc. For obtaining improved performance and mechanical properties in a wide temperature range, such components are usually made of expensive alloys, i.e., Titanium- or Nickel-based alloys. Besides, most casting dies are extremely high-priced as a result of their complicated structure and manufacturing costs.

Besides, most casting dies are extremely high-priced as a result of their complicated structure and manufacturing costs. Therefore, regular maintenance and repair are necessary to prolong their service life.Laser-assisted additive manufacturing (LAM) process for component repair has been becoming an effective candidate for part refurbishment due to its good fabricated material properties and solid bi-material bonding [1]–[3], feasibility of depositing a wide range of materials [4]–[9], high-accurate positioning and motion control using computer [10], [11], as well as easily control of processing parameters [12]–[14]. LAM can deposit material on damaged components with different materials to form functional structure [15], [16]. LAM has different names at various laboratories such as direct metal deposition (DMD), Laser Engineering Net Shaping (LENS), laser metal deposition (LMD), etc.

LAM has different names at various laboratories such as direct metal deposition (DMD), Laser Engineering Net Shaping (LENS), laser metal deposition (LMD), etc. The basic principle of such techniques is a solid freeform fabrication process which is characterized as a layer-by-layer deposition technology that can directly build near-net shape fully dense components from CAD drawings. For component repair, the laser beam302was scanned on the target damaged area to generate a melt pool while a powder stream is conjected into the melt pool. The injected power undergoes rapid melting and then solidifies on the damaged substrate when the layer jogs away. Layer by layer, the missing volume is reobtained. The shape of deposits is directly controlled by moving the substrate along a pre-defined pattern.

Layer by layer, the missing volume is reobtained. The shape of deposits is directly controlled by moving the substrate along a pre-defined pattern. It has been widely reported that LAM as a remanufacturing technique has shown great applications in automotive, aerospace and die casting industries [17]–[24].It was reported in [25], [26] that an integrated remanufacturing process generally undergoes the following steps: (1) Pre-repair inspection and decision; (2) Additive patch extraction; (3) Subtractive patch extraction; and (4) Quality Inspection. This is a universal repair routine that applied to various metallic components with diverse defects.(1) In detail, a damaged part will at first experience a complete inspection to assess the feasibility and economy of restoring it.

Since there are a variety of defects such as cracks, dents, wear, erosion, distortion, etc., and possibility in locating at a wide range of positions, the inspection and decision making is highly a case to case basis. However, the crucial factors are limited which comprise accessibility of defects, the area of damage, the geometry of the base structure and post- repair machining possibility.(2) When repair is scheduled, the repair volume of the part is then extracted to generate additive toolpath. Damage extraction is a key to guarantee a successful repair because the refilled volume is directly determined by the additive toolpath. This process usually requires creating the damaged model through 3D scanning and then conducting Boolean operation between the recreated damaged model with the nominal model.

This process usually requires creating the damaged model through 3D scanning and then conducting Boolean operation between the recreated damaged model with the nominal model. A few algorithms have been developed for regenerating the geometry of the worn-out area [11], [18], [25], [27], [28].(3) Subtractive machining succeeds the additive process since the as-deposited part mayhave unsatisfied surface roughness and inaccurate dimensions.(4) In quality inspection step, the remanufactured part is checked for any appearance of defects to ascertain that the repair is successful. In addition, the restored part is 3D scanned and then compared with the nominal model to catch possible dimensional distortion due to errors in damage extraction and heating /cooling introduced residual stress.However, the abovementioned repair routine is not complete because, in general, worn components cannot be directly deposited without pre-repair machining.

Ultrasonic cleaning or rinse can only remove surface oil or debris on worn parts but cannot remove materials surrounding inaccessible defects (such as cracks or defects in sharp corners) or contaminated materials on worn surfaces. Depositing materials on such regions cannot guarantee the interfacial bonding strength. Unsolid bonding could cause delamination when reputing the part into service. Contaminated inclusion of filler material and base decimates the mechanical properties and therefore leaves great threaten especially on key components such as aircraft turbine and compression blades. Therefore, this research introduced pre-repair machining before additive patch extraction to assure the repairability of the damaged part.303The aim of this paper is to develop pre-repair machining methodologies for component repair.

The purpose of the proposed methodologies aims to benefit repair process by not only cutting off surrounding materials around inaccessible defects but also removing contaminated layers on worn parts. Because the cut-off volume must be redeposited, the removed volume in pre- repair machining should be minimized. The overall methodologies presented in this paper take the 3D scanned model as input and based on the types of defects, different machining strategies were proposed. This study targets three types of defects: (1) surface impact defects; (2) surface erosion, corrosion and wear; (3) Cracking. The methodology is an integrated structure of 3D digitization, machining volume extraction, and pre-machining program generation.2. Pre-repair machining strategy for impact defectsThe pre-repair machining procedure for surface impact defects such as dents, notches, grooves and material overcut is illustrated in Fig.

Pre-repair machining strategy for impact defectsThe pre-repair machining procedure for surface impact defects such as dents, notches, grooves and material overcut is illustrated in Fig. 1, which includes (1) Damaged part cleaning and pre-repair inspection; (2) Model reconstruction; (3) Cut-off volume definition; (4) Target geometry acquisition; (5) Pre-machining toolpath and NC code generation; and (6) Pre-repair machining. The procedure is summarized in this section and detailed discussed in sections 2.1 – 2.5.In the beginning, the damaged part is throughout ultrasonic cleaned and inspected to determine the repair possibility. The inspection is mainly conducted by technicians due to the excess possible locations of damage. When pre-repair machining is needed, the model of the damaged part is then reconstructed. After that, the volume of material enveloping the damage is defined. Thereafter, the target geometry can be obtained.

After that, the volume of material enveloping the damage is defined. Thereafter, the target geometry can be obtained. Subsequently, based on the target geometry, pre-machining toolpath and NC node can be generated. The NC code can be exported to the machining system for machining.Fig. 1 Pre-repair machining procedure for surface impact damage2.1 Damaged part cleaning and pre-repair inspectionAn H13 tool steel block with dimensions of 50 mm × 25.4 mm × 25.4 mm was utilized as an example for illustrating pre-machining strategy of surface impact damage (Fig. 2). Ball-304indented defects were randomly prepared on the surface of the substrate using an 8 mm drill bit. The drilled holes on the block have varied depths and overlap ratios.One can see in Fig. 2 that the damaged substrate cannot be repaired directly without pre- repair machining due to at least two reasons.

2 that the damaged substrate cannot be repaired directly without pre- repair machining due to at least two reasons. One is that vertical surfaces on the side of the holes are not accessible to the laser beam and powder feed nozzle for 3-axis system. Directly depositing materials on such areas cannot guarantee the sound bi-material bonding. Therefore, materials surrounding vertical surfaces need to be machined to reveal a tilt surface to guarantee the accessibility. Another reason is that the rough tiny edges in the damaged area not only complicate the toolpath planning for material deposition but also may distort during deposition due to heat input that ruins the accuracy of deposited geometry. Therefore, it is necessary to machine off a certain amount of material enveloping the damage to provide a repairable geometry.Fig.

Therefore, it is necessary to machine off a certain amount of material enveloping the damage to provide a repairable geometry.Fig. 2 An H13 tool steel substrate with ball-indented damage on the top surface2.2 Model reconstruction of damaged component with impact damageRecreating the model of worn part is required for planning the pre-repair machining. In this research, models of damaged parts were reconstructed using a structured-light optical 3D scanner (Shining3D OptimScan-5M, Fig. 3) that can yield a validated accuracy of 20 µm. The scanner has a blue light projector that emits structured-light patterns on an object and two CCD cameras on the scanner measure the distorted dimensions of the pattern. The alteration in dimensions provides the three-dimensional coordinates of the object. In order to acquire a complete model, multiple scans are demanded for capturing different orientations of the object.

In order to acquire a complete model, multiple scans are demanded for capturing different orientations of the object. For registering point cloud from different scans to a single model, indexing targets need to be randomly pasted around the object to be scanned. Fig. 4a depicts the point cloud of the substrate and the point cloud was further processed to create the STL model shown in Fig. 4b.305Fig. 3 Shining 3D structured-light scanner for model reconstruction(a) (b)Fig. 4 Point cloud (a) and reconstructed STL model (b) of the H13 tool steel block2.3 Cut-off volume definitionThe STL model of the damaged part was sliced into a number of layers along the y-axis with a layer thickness of 0.5 mm as shown in Fig. 5a. The slicing outputs a series of layers that combined damaged and undamaged cross-sections. Damaged cross-sections also contain points in damaged and undamaged regions.

5a. The slicing outputs a series of layers that combined damaged and undamaged cross-sections. Damaged cross-sections also contain points in damaged and undamaged regions. Damaged points can be easily extracted by calculating the distance from the points to the nominal surface. A tolerance can be set so that points with distance beyond the tolerance would be defined as damaged points. Fig. 5b presents a cross-section where damage starts on the left and ends on the right.In order to machine off materials around the defects, based on the damaged cross-sections, two enveloping boundaries, U-shaped boundary (Fig. 5b) and convex-hull boundary (Fig. 7a), were utilized. Both closed profiles contain the damaged area and surrounding materials to provide a good accessibility to the LAM system.2.3.1 U-shaped boundary definitionAs shown in Fig. 5b, for U-shaped contour, two approaching lines with an inclination angle on the right) were utilized to approach the damaged cross-section.

5b, for U-shaped contour, two approaching lines with an inclination angle on the right) were utilized to approach the damaged cross-section. The θ ( damage starting point were known during the damaged point searching stage. It is necessary to find points to define the two , and 𝑆𝑆(𝑥𝑥𝑆𝑆, 𝑦𝑦𝑆𝑆, 𝑧𝑧𝑆𝑆) approaching lines. Once such points were determined, the U-shaped boundary can be obtained.on the left andand ending point𝑃𝑃𝑃𝑃𝑀𝑀𝑀𝑀𝐸𝐸(𝑥𝑥𝐸𝐸, 𝑦𝑦𝐸𝐸, 𝑧𝑧𝐸𝐸) 𝑃𝑃, 𝑃𝑃, 𝑀𝑀𝑀𝑀306and The coordinates of such points can be defined as follows: maximum z-coordinate on the damaged cross-section, which is and defined by exploring the minimum z-coordinate on the damaged cross-section, which is 𝑧𝑧𝑀𝑀 the damaged model was sliced along the y-axis, and 𝑧𝑧𝐺𝐺 randomly defined as long as 𝑥𝑥𝑁𝑁 , and the side wall inclination angle θ were defined, the coordinates according to Eq. (1).were defined by finding the were in Fig. 5b..

(1).were defined by finding the were in Fig. 5b. . Since were is much less than the minimum x-coordinate of the damaged points 𝑦𝑦𝑃𝑃 = 𝑦𝑦𝑄𝑄 = 𝑦𝑦𝑀𝑀 = 𝑦𝑦𝑁𝑁 = 𝑦𝑦𝑆𝑆 = 𝑦𝑦𝐸𝐸 and can be calculated 𝑥𝑥𝑃𝑃𝑧𝑧𝑃𝑃𝑧𝑧𝑁𝑁 𝑧𝑧𝑇𝑇𝑧𝑧𝑄𝑄 .𝑥𝑥𝑃𝑃is much larger than the maximum x-coordinate of the damaged points. Once𝑥𝑥𝑃𝑃and𝑥𝑥𝑁𝑁𝑥𝑥𝑁𝑁𝑥𝑥𝑄𝑄𝑥𝑥𝑀𝑀    x Qx M==z−z Q θ zP tan−xN+− N tanxPz M θ(1)Once the two approaching lines were defined, relationship check between the approaching and to make the approaching lines closer to the cross-section, and another checking iteration was 𝑥𝑥𝑃𝑃 conducted. Two approaching lines that firstly have intersections with the damaged curve can be 𝑥𝑥𝑁𝑁 obtained and lines before this iteration were gathered, which is shown in Fig. 5b. Finally, line andlines and the cross-section was conducted.

5b. Finally, line andlines and the cross-section was conducted. If there is no intersection, a step was applied towere added to close the U-shaped contour.𝑃𝑃𝑀𝑀 The inclination angle θ of the two approaching lines can be operator determined and adjusted flexibly according to the specifications and performance of the LAM system. It should be noted that θ has a significant effect on the cut-off volume that should be machined in the pre- repair machining step. A small θ can give a good accessibility to the LAM system but may result in a much more cut-off volume. An optimized θ needs to be determined which not only reveals good accessibility but also holds a minimized cut-off volume. For this purpose, a series of approaching lines with varied inclination angle was adopted to intersect a cross-section shown in Fig. 6a. Inclination angle θ controls the accessibility and it shouldn’t approach 90°.

6a. Inclination angle θ controls the accessibility and it shouldn’t approach 90°. A series of experiments were conducted that aim to refill slots with different sidewall inclination angle on H13 tool steel and it was found that 75° tilt angle can still yield a sound bi-material bonding and good material properties. Therefore, the range of angle θ was limited in [10°, 75°]. The area difference between the U-shaped profile and the cross-section, S∆ shown in Fig. 6b, can indicate the cut-off area. A series of approaching lines were processed on cross-sections A and B as shown in Fig. 6a-b and the relationship between S∆ and inclination angle θ was plotted in Fig. 6c.𝑃𝑃𝑀𝑀It was observed from Fig. 6c that for both cross-sections, the area S∆ decreases and then increases with the increase of angle θ. The curves indicate that neither a minimum or a maximum tile angle can result in the minimum cut-off volume. It was found that the optimal angle θ for CS A is 26° and is 31° for CS B.

The curves indicate that neither a minimum or a maximum tile angle can result in the minimum cut-off volume. It was found that the optimal angle θ for CS A is 26° and is 31° for CS B. Therefore, the optimal angle θ is highly dependent on the profile of the cross-section and should be determined for each cross-section. It can be found that for a very deep but thin damage, the minimum cut-off volume is nearly at angle 90°. This is because the closer of the angle to 90°, the less the cut-off will be.307(a)(b) Fig. 5 Model slicing and enveloping boundary determination; (a) Cross-sections of the damaged block (layer thickness = 0.5 mm); (b) U-shaped enveloping boundary;(a)(b)308(c) Fig. 6 Sidewall inclination angle optimization for the U-shaped boundary.

6 Sidewall inclination angle optimization for the U-shaped boundary. (a) U-shaped boundary with varied approaching lines; (b) U-shaped boundary for cross-sections A and B; (c) Relationship between area difference S∆ and sidewall inclination angle θ2.3.2 Convex-hull boundary definitionOne can see in Fig. 5b that for U-shaped boundary, the bottom connecting line QM is parallel to the x-axis, and this may result in extra material cut-off in the polygon GMETG . This material over-cut comes worse when one deep defect is presented while majority defects are shallow. To further minimize cut-off volume, the convex-hull boundary was obtained as shown in Fig. 7a. The convex-hull of any polygon can be easily obtained based on existing algorithm [29]. However, one should notice that the convex-hull cannot be directly utilized as the boundary for pre-machining owing to the possibility of the steep area as shown in Fig. 7a.

However, one should notice that the convex-hull cannot be directly utilized as the boundary for pre-machining owing to the possibility of the steep area as shown in Fig. 7a. The line segments in those regions have big tangents that cannot guarantee the accessibility. Therefore, such lines need to be tilted. The algorithm for tilting lines with big tangent is schematically depicted in Fig. 7b and discussed below.Suppose polygonis a convex-hull of one cross-section. The inclination needs to be can be calculated according to Eq. 2, and γ is the complementary angle of θ. The ′ ′ ′ , 𝑦𝑦𝑖𝑖 𝑃𝑃𝑖𝑖 (𝑥𝑥𝑖𝑖 also needs to be also exceeds the allowed angle and therefore, 𝑃𝑃𝑖𝑖+1 . However, the . After that, the shifted polygon is 𝑃𝑃𝑖𝑖+1 needs to be further ′ ′ ′ 𝑃𝑃𝑖𝑖+2 𝑃𝑃𝑖𝑖+7 ⋯ 𝑃𝑃𝑖𝑖+6 𝑃𝑃𝑖𝑖 . After the second iteration, the shifted polygon became ′ ′ ′ 𝑃𝑃𝑖𝑖+1 𝑃𝑃𝑖𝑖 𝑃𝑃𝑖𝑖 . Finally, the contour.

After the second iteration, the shifted polygon became ′ ′ ′ 𝑃𝑃𝑖𝑖+1 𝑃𝑃𝑖𝑖 𝑃𝑃𝑖𝑖 . Finally, the contour . There is another iteration to move ′ 𝑃𝑃𝑖𝑖+1 is obtained that satisfies the accessibility requirement, and therefore, can be ′′ 𝑃𝑃𝑖𝑖+7 ′′′ 𝑃𝑃𝑖𝑖+7angle of rotated around 𝑃𝑃𝑖𝑖𝑃𝑃𝑖𝑖+1 where 𝑃𝑃𝑖𝑖+1 inclination angle of rotated around inclination angle rotated aroundis indicated as θ. Because θ is beyond the allowed accessible angle,𝑃𝑃𝑖𝑖𝑃𝑃𝑖𝑖+1 ⋯ 𝑃𝑃𝑖𝑖+6𝑃𝑃𝑖𝑖+7 . The coordinate of is the coordinate of ′ 𝑃𝑃𝑖𝑖 (𝑥𝑥𝑖𝑖+1, 𝑦𝑦𝑖𝑖+1, 𝑧𝑧𝑖𝑖+1) to 𝑃𝑃𝑖𝑖+1𝑃𝑃𝑖𝑖+2 still surpasses the desired angle and therefore, ′ 𝑃𝑃𝑖𝑖+1 toto𝑃𝑃𝑖𝑖′ , 𝑧𝑧𝑖𝑖)′ 𝑃𝑃𝑖𝑖+1 to′′ 𝑃𝑃𝑖𝑖′′ 𝑃𝑃𝑖𝑖 used as a boundary for cut-off volume definition. 𝑃𝑃𝑖𝑖′′ 𝑃𝑃𝑖𝑖+1 ′′′ ′′′ 𝑃𝑃𝑖𝑖+1′′ ⋯ 𝑃𝑃𝑖𝑖+6 ′′′ ⋯ 𝑃𝑃𝑖𝑖+6′′ 𝑃𝑃𝑖𝑖+7′′′ 𝑃𝑃𝑖𝑖+7309The aforementioned algorithm was applied to a cross-section as shown in Fig. 7a and the optimized contour is depicted in Fig. 7c. As illustrated in Fig.

7a and the optimized contour is depicted in Fig. 7c. As illustrated in Fig. 7c, the steep area in the original convex hull was successfully tilted.(a)(b)(c) Fig. 7 Convex-hull boundary definition. (a) Convex-hull processed on a cross-section; (b) Sidewall inclination angle optimization for convex-hull boundary; (c) Optimized contourwheny i+ 1<y i(Decreasing) :wheny i+ 1>y i(Increasing) :x i'=x i+ 1−  − y y + 1 i i γ cos  2−(y i+ 1−y i)2x i+ 1'=x i+  − y y + 1 i i γ cos  2−(y i+ 1−y i)2(2)y i'=y iy i'=y iz i'=z iz i'=z i3102.4 Target geometry acquisitionU-shaped and convex-hull strategies were processed on the STL model with ball- ° and indentations (Fig. 8a). For U-shaped strategy, side wall with fixed inclination angle optimized angle were both conducted, and the target models were shown in Figs. 8b and 8c, respectively.

8a). For U-shaped strategy, side wall with fixed inclination angle optimized angle were both conducted, and the target models were shown in Figs. 8b and 8c, respectively. The model processed using convex-hull strategy was presented in Fig. 8d.45θ =The volume of the untreated damaged model (Fig. 8a) and volumes of processed models (Figs. 8b-d) using proposed strategies were summarized in Table 1. The volumes of machined models are different because the strategies for target geometry acquisition are different. It can be observed from Table 1 that the U-shaped strategy with optimized sidewall inclination angle will machine off less material compared with non-optimized U-shaped strategy. However, convex-hull strategy results in the least amount of cut-off volume, which is expected. Therefore, in the next machining step, convex-hull strategy-generated geometry was adopted as the guide model for machining toolpath and NC code generation.(a) (b)(c) (d)Fig.

Therefore, in the next machining step, convex-hull strategy-generated geometry was adopted as the guide model for machining toolpath and NC code generation.(a) (b)(c) (d)Fig. 8 Untreated and processed models; (a) Model of the damaged substrate; Model pre- machined using U-shaped strategy (b), U-shaped strategy with optimized sidewall inclination angle (c) and convex-hull strategy (d)Table 1 Volume of the model with varied pre-machining strategiesModelVolume (mm3) Cut-off volume (mm3)Damaged22245.1U-shaped machined21216.31028.8U-shaped machined with optimized sidewall angle 21251.6993.5Convex-hull machined21331.8913.33112.5 Pre-repair machining program generation and pre-machiningOnce the target geometry was obtained, it was loaded to Mastercam software to generate the machining toolpath (Fig. 9a) and program. The milling parameters are listed in Table 2.

9a) and program. The milling parameters are listed in Table 2. The damaged substrate was clamped on a fixture and machined on a Fryer MC-30 with the setup shown in Fig. 9b. The machined part is shown in Fig. 9c.Table 2 Pre-repair machining parameters for surface impact damageToolFeed rateSpindle speedStepdownStepover percentageToolpath type¼’’ ball-end mill508 in/min2000 r/min0.533 %High-speed spiral(a) (b) (c)Fig. 9 Pre-machining toolpath (a), experiment setup (b) and machined substrate (c)3. Pre-repair machining strategy for surface erosion, wear and corrosionCasting dies are subjected to rapid heating and cooling cycles during their service that causes dimensional distortion and cracking. In addition, frequently contacting with casting alloys such as liquid aluminum alloys causes surface defects such as erosion, wear, and corrosion. It is reported that thermal fatigue and erosion are major contributions to the failure of casting dies [30].

It is reported that thermal fatigue and erosion are major contributions to the failure of casting dies [30]. Die surface flaws destruct the quality of casted copies, therefore needs frequent scheduled repair.Directly depositing materials on damaged die surfaces is not suggested because of the following reasons. Firstly, damaged surface usually contains impurities such as oxidized and contaminated materials. Surface pitting may also cause numerous tiny localized defects. Coating materials on such untreated surfaces cannot guarantee the solid bi-material bonding which may cause unpredictably delamination during application. Secondly, LAM requires a certain surface roughness for material deposition (shiny surface requires high laser power and rough surface requires lower laser power). Damaged surfaces may not have a constant, defined roughness, whereas a machined surface can provide this.

Damaged surfaces may not have a constant, defined roughness, whereas a machined surface can provide this. Fortunately, such surface defects are usually superficial and can be easily removed by cutting off a thin layer of material from the target surface. The thickness of machined material can be determined as long as the machined surface is defect- free. Once such layer is machined, material can be coated on such region to reobtain the designed geometry.312The current research proposes a general procedure for treating surfaces with erosion, wear, and corrosion. In general, the damaged component is 3D scanned to recreate its STL model. After that, the model is loaded to CAM software, and the target regions are selected. Machining parameters including cut-off thickness are determined and machining program is generated. Subsequently, the program is transferred to CNC for pre-machining.

Machining parameters including cut-off thickness are determined and machining program is generated. Subsequently, the program is transferred to CNC for pre-machining. The following sections detailed the pre-repair machining procedure for such damage.3.1 Model reconstruction of a casting die with surface defectsA casting die shown in Fig. 10a has damaged regions on its working convex area that dominated by tiny rough surfaces. In order to machine off a thin layer of material from the part, the STL model of the die was generated using the configuration in Fig. 3 and the recreated model is shown in Fig. 10b. It should be pointed out that creating the model of the whole part is not necessary for pre-repair machining. This is because only the surface with defects needs to be targeted and other regions will not be treated.

This is because only the surface with defects needs to be targeted and other regions will not be treated. Generating the whole model not only dramatically slow down the 3D scanning process, especially for a large complex structure, but also sophisticate the toolpath generating process due to the large size of the imported model.(a) (b)Fig. 10 (a) A casting die with surface erosion, wear, and corrosion; (b) STL model of the damaged region of the casting die3.2 Machining volume determination, program generation, and pre-machiningThe volume of materials that need machining depends on the depth of surface defects. The defects of surface erosion, wear and corrosion are usually superficial and can be effectively removed through one pass machining. The cut depth for machining can be determined based on the condition of damaged components. For the casting die in Fig. 10a, materials with 0.5 mm layer thickness were machined. The toolpath was generated as shown in Fig.

For the casting die in Fig. 10a, materials with 0.5 mm layer thickness were machined. The toolpath was generated as shown in Fig. 11a and the simulated machined part is shown in Fig. 11b.313(a) (b)Fig. 11 (a) Pre-machining tool path generation; (b) Pre-machined model of the die4. Pre-repair machining strategy for crackingThe heating and cooling cycles during metal casting cause cyclic compressive and tensile stress conditions which led to thermal fatigue to the casting die [31]. That is why H13 hot work tool steel is usually adopted as die material owing to its high hardenability and good thermal fatigue resistance. Thermal fatigue cracking is a common failure in die-casting dies and engine blades after thousands of shots.

Thermal fatigue cracking is a common failure in die-casting dies and engine blades after thousands of shots. The cracking is first to like to appear at corners and edges with small radius [32].Detailed machining procedure for cracking highly depends on the appearance of cracks and can be determined only after considering a number of factors, such as the depth and length of cracks, surrounding structures, accessibility of cracks to machining tool and LAM system, surface or internal cracking, etc. Therefore, this paper presents some candidate machining approaches that can be considered for cracking removal.Table 4 lists a few machining methods for cracking. Cracks are classified into shallow and deep cracks because they may need different machining methods. For example, for shallow cracks, they can be easily and effectively removed by surface grinding. Therefore, 3D scanned models of such parts are not needed.

For example, for shallow cracks, they can be easily and effectively removed by surface grinding. Therefore, 3D scanned models of such parts are not needed. For deep cracks, Electro discharge machining (EDM) and CNC (computerized numerical control) machining can be utilized. EDM can cut off materials with cracks much more efficiently, but it is only applicable to cracks that can be accessible to EDM wire. If the cracking is surrounded by other structures, EDM is not applicable and therefore, CNC machining can be considered. Such machining procedures will consume more material over surface grinding. One should know that internal cracking is not easy to observe visually and then may require nondestructive testing. Once the internal cracking is located, a slot can be made to remove the cracks. However, machining a slot is not always working because a very deep slot is difficult to refill, especially when the edge of the slot is near vertical.

However, machining a slot is not always working because a very deep slot is difficult to refill, especially when the edge of the slot is near vertical. In such cases, the U-shaped strategy can be applied or the portion of the component with cracks can be completely cut-off and rebuild. This approach is applicable to any cracks regardless of depth, but it will cut-off a huge amount of material. Therefore, one needs to evaluate the cost by repairing them over replacement314with a new copy. In this paper, a titanium engine blade with cracks was machined using EDM to show the machining procedure.Table 4 Candidate machining approaches for cracking4.1 Pre-repair machining strategy for crackingAn engine blade was utilized to demonstrate the pre-repair machining strategy for cracking. Trailing edge cracking is a common failure in turbine or compressor blade mainly due to excessive stress, overloading, overheating or defective materials.

Trailing edge cracking is a common failure in turbine or compressor blade mainly due to excessive stress, overloading, overheating or defective materials. Several studies have been targeting on blade repair, mainly focusing on damage extraction [18], [33], [34], material deposition process [35]–[37]and repair automation [38], [39], while no reporting on strategies of pre-machining. It is required to cut off the material surrounding cracks because non-machined cracks give no accessibility to LAM system, as shown in Fig. 12a. This is why the defects in blade edge reported in [19], [38], [40] have U or V-shaped geometry.To perform pre-machining, the blade as shown in Fig.12a was digitally scanned to acquire its 3D model as shown in Fig. 12b. After that, the defective area was selected as shown in Fig. 12c. By selecting the damaged area, the points located at the worn domain can be easily extracted which is shown in Fig.

12b. After that, the defective area was selected as shown in Fig. 12c. By selecting the damaged area, the points located at the worn domain can be easily extracted which is shown in Fig. 13a, and subsequently, the convex hull of the point set was obtained. It should be pointed out that the convex hull cannot be directly used for pre-machining because the bottom portion of the convex hull blocks the accessibility. In order to create an open V-shaped geometry, the line segments of the convex hull were tilted according to the algorithm illustrated in 2.3.2. The maximum sidewall inclination angle θ is limited at 60°. The finally optimized contour was shown in Fig. 13b. EDM wire then follows the optimized contour for pre-machining and the pre-machined blade is shown in Fig. 13c.315(a) (b) (c)Fig. 12 (a) Titanium blade with cracks; (b) Reconstructed model of the blade; (c) Selection of points located at the damaged area(a) (b) (c)Fig.

13c.315(a) (b) (c)Fig. 12 (a) Titanium blade with cracks; (b) Reconstructed model of the blade; (c) Selection of points located at the damaged area(a) (b) (c)Fig. 13 (a) Damaged point set extraction and its convex hull; (b) Optimized contour; (c) Blade after pre-machining5. ConclusionPre-repair machining of damaged components is crucial before additive manufacturing to make sure the damage is accessible and the substrate is free of contamination. For this purpose, this paper presents several pre-machining strategies targeting at surface defects and cracking. Damaged part was first inspected and machining tool was determined. Reconstructed model of the damage component is usually required for pre-repair machining toolpath generation. For surface impact defects, two strategies, U-shaped, and Convex-hull strategies were introduced for cut-off volume definition. The convex-hull method results in less material removal and is, therefore, preferred.

The convex-hull method results in less material removal and is, therefore, preferred. For surface erosion, wear, and corrosion, a thin layer of material at the damaged region was removed to get rid of contaminated layers or numerous tiny defects. Approach for removing cracking should be determined after considering several factors such as the depth and length of cracks, surrounding structures, accessibility of cracks, surface or internal cracking, etc. This study presents an approach for removing cracks in blade edge using wire-EDM.316AcknowledgmentThis project was supported by TOYOTA/Bodine Aluminum, National Science Foundation Grants CMMI-1547042 and CMMI-1625736, and the Intelligent System Center, Center for Aerospace Manufacturing Technology, and Material Research Center at Missouri S&T. Their financial support is greatly appreciated.Reference[1] R. Raju, M. Duraiselvam, V. Petley, S. Verma, and R.

viewed PaperDEVELOPMENT OF PROCESSING PARAMETERS FOR THE SELECTIVE LASER SINTERING OF CARBON FIBER REINFORCED POLYPHENYLENE SULFIDE WITH A TOOLING APPLICATIONScott E. Snarr*, Patrick L. Snarr*, Joseph Beaman Jr.*Department of Mechanical Engineering, The University of Texas at Austin, Austin, TX 78712AbstractThe ongoing development of processing parameters for advanced thermoplastic materials to be fabricated via selective laser sintering (SLS) is rapidly advancing the potential industrial applications of the manufacturing method. This research focuses on the development of SLS processing parameters and a tooling application for carbon fiber reinforced Polyphenylene Sulfide (CF-PPS), a composite material that is novel to SLS. A high temperature SLS research machine was used to identify suitable processing parameters for the material along with the tensile strength and geometrical accuracy associated with those parameters.

A high temperature SLS research machine was used to identify suitable processing parameters for the material along with the tensile strength and geometrical accuracy associated with those parameters. Utilizing the previously identified parameters, a repair mold for an electronic cable assembly was fabricated. Mechanical tests were performed on fabricated CF-PPS parts to evaluate the performance of the material under the mold’s normal operating conditions. The additively manufactured CF-PPS mold was deemed viable for production and was shipped to our sponsor for further evaluation.Introductionis an additive manufacturing process capable of fabricating fully dense polymer parts with complex geometries that cannot be produced by other manufacturing methods [1] [2]. The process begins with a counter-rotating roller spreading a thin layer of powder over a piston. A laser then scans the cross-sectional area of the desired part(s).

The process begins with a counter-rotating roller spreading a thin layer of powder over a piston. A laser then scans the cross-sectional area of the desired part(s). This raises the pre-heated polymer powder above its melting temperature, subsequently fusing the powder material together upon cooling. The piston then moves down a single layer and another layer of powder is spread on top of the previously sintered material. The process then repeats itself, layer-by-layer, to build up a three dimensional part(s) [3]. The SLS manufacturing low cost parts for low volume, high method excels in quickly delivering high efficiency, complexity component production. Seeking these benefits, the medical, aerospace, and tooling industries have been rapidly expanding their applications of this developing technology.Selective Laser Sintering (SLS)The ongoing development of processing parameters for advanced thermoplastic materials is also broadening the potential applications of the SLS process.

This research focuses on the industry applications for carbon fiber development of processing parameters and potential to SLS. reinforced polyphenylene sulfide (CF-PPS), a composite material Polyphenylene sulfide (PPS) is a high performance thermoplastic that offers exceptional mechanical strength, chemical resistance, and high operating temperatures [4]. The addition of carbon fibers to PPS enhances both the stiffness and strength to weight ratio of the material while also elevating the material’s heat distortion temperature. The combination of these qualities make it a desirable material to be used in tooling applications. An unnamed sponsor desired to develop processing parameters for CF-PPS in order to fabricate a mold for repairing an electronics cablethatis novel206assembly in adverse situations and conditions.

The cable assembly has an exterior plastic casing that is in place to protect the electronics cable inside, but the casing frequently deteriorates causing the cable to be exposed. The fabricated mold would allow for the exposed electronics cable to be placed inside and a new protective plastic casing to be molded over the cable assembly. This results in significant cost savings as well as reduced lead times when compared to replacing the damaged cable assembly. Through this research, SLS processing parameters for building with CF-PPS were developed and the electronics cable assembly repair mold was fabricated. After construction, the mold was shipped to our sponsor and put through a series of tests to evaluate its performance. The fabricated quality of the mold and initial testing results showed promise for the use of CF-PPS in future tooling applications.Experimental Methods and ProceduresParameter Developmentto the SLS process is parameter development.

Parameter development refers to identifying the appropriate processing window and set point for each of the important variables in the SLS process. These variables include laser power, laser scan speed, layer thickness, hatch spacing, and powder bed preheat temperature. If any of these variables are outside of the necessary processing window, the novel material will not successfully build in SLS. The forthcoming paragraphs will introduce a proven method for parameter development for a novel material to be used in SLS.The first step when introducing a novel materialThe work outlined in this paper was performed on the Laser Additive Manufacturing Pilot System (LAMPS) at the University of Texas at Austin. LAMPS is a research grade SLS machine capable of processing thermoplastic powders with melting temperatures approaching 400 ℃. This is made possible by an insulated upper build chamber that houses three quartz lamps directly over the powder bed to preheat to the desired temperature.

This is made possible by an insulated upper build chamber that houses three quartz lamps directly over the powder bed to preheat to the desired temperature. Two of the quartz lamps are directed at the powder bed, and the remaining lamp is pointed at the powder drop to help preheat the new powder layer prior to spreading. Preheating the powder in the SLS process is important in maintaining a uniform powder bed temperature to help fight part curling during the build process. LAMPS is also outfitted with two thermal cameras and a visual camera to provide real time monitoring of the build taking place. The data generated from the two thermal cameras is used to monitor and control the thermal environment during operation and also for research purposes to help characterize the SLS process. The visual camera can be monitored by the operator to have eyes inside of the build chamber and to ensure everything looks as it should throughout the SLS build process.

The visual camera can be monitored by the operator to have eyes inside of the build chamber and to ensure everything looks as it should throughout the SLS build process. On the next page, Figure 1 shows a CAD rendering of the LAMPS research SLS machine at UT Austin.207Figure 1: CAD rendering of the LAMPS SLS machineLAMPS utilizes a 40 W CO2 laser and a galvo mirror system to scan two-dimensional cross sections of the powder bed. The input power and scan speed of the laser on the powder bed directly correlates to the energy input into the material. When selecting which laser power to use in LAMPS, a percentage of the total power is the input. In theory, 50% power would result in 20 W of power being output from the laser and into the powder bed. In practice, however, this is not the case as there are several factors that reduce the power of the laser.

In practice, however, this is not the case as there are several factors that reduce the power of the laser. These factors include inefficiencies in the optic track of the laser, and a degradation of the laser over time, contaminated laser window inside the build chamber. To account for this, the laser is tested before and after each build to quantify the laser power delivered to the powder bed during the build. This is important because during longer duration builds, the laser window can become contaminated with powder or off-gassing from the material and reduce the laser power seen by the powder bed. This issue is somewhat mitigated by flowing nitrogen over the window but cannot be completely eliminated. Figure 2 shows the laser power curve obtained before the tensile bar builds from this study.

This issue is somewhat mitigated by flowing nitrogen over the window but cannot be completely eliminated. Figure 2 shows the laser power curve obtained before the tensile bar builds from this study. An after measurement is not presented as no contamination was detected due to the short duration of the builds.Figure 2: Measured laser power at build surface208This research utilized a modified version of the processing parameter developmentprocedure laid out by Snarr et al. [5]. The modified approach contained the following steps:1. Verify the material’s powder particle size is suitable for the SLS process 2. Identify sintering window for the material 3. Use visual inspection to determine powder bed preheat temperature 4. Sinter energy density grid to identify laser scanning parameters 5.

Identify sintering window for the material 3. Use visual inspection to determine powder bed preheat temperature 4. Sinter energy density grid to identify laser scanning parameters 5. Fabricate tensile bars to quantify part strength and geometrical accuracy for parameters A modified version of the parameter development steps was used because the end goal of this research was not to optimize parameters for strength or geometrical accuracy, but to fabricate a functioning repair mold for the sponsor.The first two steps in the parameter development process involved characterization of the CF-PPS powder to determine its suitability for the SLS process. This included checking the average mean diameter and sphericity of the powder particles along with identification of the material’s glass transition and melting temperatures.

This included checking the average mean diameter and sphericity of the powder particles along with identification of the material’s glass transition and melting temperatures. A Scanning Electron Microscope (SEM) was used to image the powder and verify the powder was composed of high sphericity particles, which facilitate smooth spreading during the SLS process. Collected images were then put through an ImageJ software analysis where particle cross-sectional area measurements were recorded for representative samples of the powder. The average particle diameter and a particle diameter distribution for the CF-PPS powder were then calculated from this data. The particles were sufficiently spherical and their mean particle diameter was within the desired range for use in the SLS process (approximately 30 - 80 um). Once this analysis was complete, the powder was then tested using a Differential Scanning Calorimetry (DSC) machine.

Once this analysis was complete, the powder was then tested using a Differential Scanning Calorimetry (DSC) machine. The DSC machine runs the powder through a heating and cooling cycle in order to identify the material’s glass transition and melting temperatures. The region between these two temperatures is known as the material’s sintering window. It is critical to the SLS process as this is the temperature range to which the material is preheated before it is sintered. The DSC process clearly identified the desired temperatures and subsequently the sintering window. At the request of our sponsor, none of the results or information collected in these steps is presented due to the proprietary nature of the material’s composition.Once the powder was deemed suitable for the SLS process and a sintering window identified, the material was put into the LAMPS machine to identify the preheat temperature. The shimmer test is a commonly used method to identify this set point.

The shimmer test is a commonly used method to identify this set point. A uniform layer of powder is spread in the machine and then the preheat temperature is slowly increased until the powder bed begins to slightly shimmer. Once the powder begins to shimmer, melting has started to occur and the upper limit of the powder preheat has been reached. The preheat temperature is then set several degrees below this upper limit in order to require minimal energy input from the laser to melt the material during the sintering phase.209The next step in the parameter development process was to utilize a parameter grid to identify laser parameters for building in the CF-PPS material. The standard volumetric energy density equation, commonly used in additive manufacturing applications with continuous wave lasers, is presented below in Equation 1.

The standard volumetric energy density equation, commonly used in additive manufacturing applications with continuous wave lasers, is presented below in Equation 1. This equation lumps the standard laser and machine parameters used in the sintering processes together to create a single value representing the total volumetric energy density ( ) applied to the material. The variables and their units used in the equation are as follows: for hatch spacing in mm, and for layer thickness in mm [6]. 𝑧𝐸𝑃for laser power in watts,𝑣for scan speed in mm/s,ℎ𝐸 =𝑃 𝑣 * ℎ *𝑧 [𝐽𝑚𝑚3 ]Several parameters had already been set based on the laser spot size or the machine itself. For example, the layer thickness for the SLS machine was .1 mm (100 um) and the hatch spacing was selected as .3 mm (300 um) based on the laser spot size of the machine (approximately .5 mm or 500 um) and standard procedures.

In order to identify the two remaining laser parameters, laser power and scan speed, parameter grids were used. These grids are used as a visual inspection tool to look for complete sintering of the powder bed and narrow down a parameter range. The color of the sintered material and slight sinking of the material into the powder bed are indicators used to find the most promising parameter set. A few different parameter grids were run with varying ranges of energy densities. The final sintered parameter grid is shown on the next page in Figure 3. In this grid, each row has a different laser power. Starting from the top row with a command of 50% laser power (6.9 W), each row increased by 10% laser power. The bottom row had a laser power command of 90%, resulting in 14.1 W of laser power. Each vertical column had a different laser scan speed. Starting from the left with a scan speed of 1500 mm/s, each row decreased by 200 mm/s leaving the far right row with a scan speed of 700 mm/s.

Each vertical column had a different laser scan speed. Starting from the left with a scan speed of 1500 mm/s, each row decreased by 200 mm/s leaving the far right row with a scan speed of 700 mm/s. The bracketed square identifies the selected parameter set to move forward with. It was chosen because it showed signs of achieving full melting (dark color change and slight sinking into the powder bed) with no loss of geometrical accuracy from overmelting or degradation of the powder. This final set of selected parameters was used for the remainder of the experiments to fabricate tensile bars and the electronics mold. The final parameters are presented on the next page in Table 1.210EQ.

The final parameters are presented on the next page in Table 1.210EQ. 1 [6]Figure 3: Parameter grid used to identify final scanning parametersTable 1: Final parameters used for tensile bar and mold fabricationLaser Power Scan Speed Hatch Spacing Layer Thickness Laser Spot Size Preheat Temperature14.1 W900 mm/s300 um100 um500 um280 °CWith the final parameter set identified, the last step in the parameter development process was to fabricate several tensile bars in order to quantify the tensile strength, compression strength, and geometrical accuracy for the parameters. Two builds, each containing seven standard ASTM Type IV tensile bars, were completed to construct a total of 14 tensile bars. On the next page Figure 4 shows the sintered tensile bars during the build process. The testing procedures used to collect data for these tensile bars are described in the following sections.

On the next page Figure 4 shows the sintered tensile bars during the build process. The testing procedures used to collect data for these tensile bars are described in the following sections. This study did not seek to optimize for strength or geometrical accuracy, but instead fabricated the tensile specimen to quickly quantify some of these values as a base case for the mold fabrication.Once the tensile specimen had been fabricated, they were taken out of the build chamber and the parts were inspected to ensure they were fairly dense and had reasonable geometrical accuracy. The as-built tensile bars can be seen on the next page in Figure 5. They were deemed acceptable and allowed the study to continue towards its main goal, fabrication of the electronics repair mold. The same set of parameters were then used to build the full-scale repair mold.

The same set of parameters were then used to build the full-scale repair mold. The outcome from the fabrication of the mold is presented in the subsequent results section.211Figure 4: Sintered tensile bars in the build chamberFigure 5: As-built tensile bars212Testing ProceduresSeveral tests were performed on samples fabricated from the novel CF-PPS material to begin validation for the use of the material for the repair mold application. These tests include tensile testing, geometric accuracy measurements, and compression testing. The results of these tests will provide insight is a candidate for the desired application.into whether this novel materialTensile testing was performed on fabricated CF-PPS tensile specimens in accordance with ASTM D638: Standard Test Method for Tensile Properties of Plastics [7]. These tests were performed on an Instron 3345 with a 5 kN load cell. ASTM Type IV tensile specimens were fabricated to complete the testing.

These tests were performed on an Instron 3345 with a 5 kN load cell. ASTM Type IV tensile specimens were fabricated to complete the testing. Figure 6 shows the nominal dimensions for a Type IV tensile specimen. A total of 5 tensile specimens were tested at a crosshead velocity of 5 mm/min as required by the standard.Figure 6: ASTM Type IV Tensile Specimen (units in mm)To test the geometrical accuracy achieved with the selected parameter set, the fabricated tensile specimens were measured at three locations using digital calipers and compared to the nominal dimension of the tensile specimen. Three measurements were taken to quantify the accuracy of the tensile specimens in the X and Y directions of the build. These measurements include the overall length (LO), the overall width (WO), and the width of the narrow section (W).

These measurements include the overall length (LO), the overall width (WO), and the width of the narrow section (W). The overall length tests the accuracy of the print in the x-axis, while the overall width and width of the narrow section reveal accuracy information of the y-axis.The final test that was performed was a standard compression test on an MTS Criterion Model 45 with a 100 kN load cell. This test was performed to begin understanding how the CF-PPS mold would handle the applied compression force when being used to repair the electronic cable assemblies. In use, the mold will be at an elevated temperature while the compression is applied, however, timely access to a machine with this capability was not accessible. To perform the compression test, a thin rectangular cross section of the CF-PPS was placed between the two compression heads and the applied force was set to a max of 95 kN.

To perform the compression test, a thin rectangular cross section of the CF-PPS was placed between the two compression heads and the applied force was set to a max of 95 kN. The crosshead was moved at a rate of .01 mm/s.213ResultsTensile Testing ResultsFigure 7: Stress-Strain curve for CF-PPS fabricated via SLSThe tensile test results provided several insights into CF-PPS parts made using the SLS process. First, the bars demonstrated extremely brittle behavior having little to no ductility. No visible necking of the bars occurred during the tensile tests. Second, the average ultimate tensile strength of the CF-PPS was measured to be just under 32 MPa. This was much lower than expected for CF-PPS (previously tested PPS parts averaged near 43 MPa) but this shortcoming can be attributed to several things. The scanning parameters developed in this study were not optimized for strength in any way. Further exploration of scanning parameters could increase the strength.

The scanning parameters developed in this study were not optimized for strength in any way. Further exploration of scanning parameters could increase the strength. Also, when breaking out the bars from the builds, the powder bed was very loose, showing no signs of being near the material’s melting point. This indicates the preheat temperature could be significantly increased in the future which plays a critical role in creating fully dense and stronger parts. Furthermore, the tensile bars had slight shifting in the Y direction which could potentially cause misalignment of the tensile testing grips, introducing a bending moment during the tensile test and ultimately reducing the ultimate tensile strength. Much higher strengths could be achieved in the future by addressing some of these key points. One additional insight taken from the stress-strain curve was the average modulus of elasticity of the CF-PPS samples, which was calculated to be 4.66 GPa.

One additional insight taken from the stress-strain curve was the average modulus of elasticity of the CF-PPS samples, which was calculated to be 4.66 GPa. The carbon fiber reinforcement significantly stiffens the polymer resulting in an increased modulus of elasticity when compared to previously built PPS parts with a modulus of approximately 2.83 GPa.214Geometrical Accuracy ResultsThe results from the geometrical accuracy tests are housed in Table 2. The nominal dimensions for the overall length, overall width, and width of the narrow section are 115 mm, 19 mm, and 6 mm, respectively for the ASTM Type IV tensile specimen.

The nominal dimensions for the overall length, overall width, and width of the narrow section are 115 mm, 19 mm, and 6 mm, respectively for the ASTM Type IV tensile specimen. The measurements from the four tensile specimens were averaged, and then the percentage error from nominal was calculated.Table 2: Results from geometrical accuracy testingOverall Length (LO)Overall Width (WO)Width of Narrow Section (W)Nominal [mm]115196Average [mm]112.0921.777.15Standard Deviation [mm]0.170.110.16Percent Error [%]2.5314.5719.08It can be observed in the results of the geometrical accuracy testing that the tensile specimens were much closer to the nominal dimension in the X direction than they were in the Y direction. This can be directly attributed to shifting in the Y direction that was observed in the final tensile specimens. If a new layer is shifted slightly compared to the layer below it, this will cause oversized dimensions in the Y direction.

If a new layer is shifted slightly compared to the layer below it, this will cause oversized dimensions in the Y direction. This phenomenon was observed in the overall width, and width of the narrow section, the two dimensions controlled by the Y direction accuracy. This is the same phenomenon that was discussed previously that led to lower tensile strengths than previously tested samples. The results show that further optimization is needed in both the X and Y direction to allow for manufacturing of tight tolerances.Compression Test ResultsThe compression test was a quick look at how the CF-PPS material would hold up when subjected to a compression force similar to how the electronics mold will be used. The electronics mold will be clamped together with a force of just below 9 kN. With a surface area of approximately 10,000 square millimeters, the mold will only be subjected to about 0.9 MPa of stress.

With a surface area of approximately 10,000 square millimeters, the mold will only be subjected to about 0.9 MPa of stress. In the compression test, the CF-PPS part was loaded up to 95 kN corresponding to a stress of around 190 MPa in the part tested. The part survived the compression and showed no signs of cracking or failure. There was some deformation of the part but most of this is believed to be due to densification from the part’s porosity. This test provided evidence that the electronics mold will be able to handle the compression force even at elevated temperatures as the CF-PPS piece showed no signs of failure at a stress over 200 times the stress it will be subjected to during operation.215Mold Application PrototypingFigure 8: Completed electronics moldThe development of the processing parameters detailed throughout this paper were all in an effort to fabricate the electronics mold that the sponsor desired.

Figure 8 shows the finished mold that was built using the previously determined parameter set. The mold turned out very good with sharp geometrical accuracy and no issues with the shifting that the tensile bars had. The only issue with the mold was the male mating pegs, seen in the top right corner of the figure, were oversized. This was contributed to a file issue and they were subsequently sanded down to fit into the female counterpart. The mold was packaged and shipped to the sponsor for it to undergo testing for use in the field. Early results from this testing are promising and validate the potential use of additively manufactured CF-PPS for a multitude of tooling applications.216ConclusionThe purpose of this research study was to develop selective laser sintering parameters for a novel CF-PPS material in order to fabricate a repair mold for an electronics cable assembly.

In order to develop parameters, first, a series of tests were performed to determine the suitability of the material for the SLS process. Next, important scanning parameters were identified using a shimmer test and parameter grids. Once parameters were selected, tensile bars were fabricated in order to quantify several performance aspects of the material. Finally, the repair mold was built using the identified parameters and shipped to our sponsor to undergo further testing.

Finally, the repair mold was built using the identified parameters and shipped to our sponsor to undergo further testing. The following results summarize the findings of this research:A set of parameters were developed for fabricating CF-PPS parts utilizing the selective laser sintering process (Summarized in Table 1)Several properties of the CF-PPS parts were quantified:Ultimate tensile strength - 32 MPa ○ Modulus of elasticity - 4.66 GPa ○ Geometrical accuracy summarized in Table 2The CF-PPS material withstood compression stresses 200 times larger than the expected operating conditions of the moldA usable repair mold was fabricated and shipped to the sponsor for further testing217References[1] Goodridge, R. D., Tuck, C. J., & Hague, R. J. M. (2012). Laser sintering of polyamides andother polymers. Progress in Materials science, 57(2), 229-267.[2] Hague, R., Dickens, P., & Hopkinson, N. (Eds.). (2006).

velopment of Reverse Engineering System for Machine EngineeringUsing 3D Bit-map DataTatsuro Yashiki* and Tarou Takagi*Power & Industrial Systems R&D Laboratory, Hitachi, Ltd.AbstractIn this paper, the authors describe a reverse engineering system, named BitCAT, for machineengineering using 3D bit-map data obtained from CT digitizers. BitCAT can compose B-Rep CADmodels by fitting parametric primitives, such as flat planes and cylinders, onto 3D bit-map surfaces. To find boundaries of the primitives effectively, a novel interactive “growing surface method” is developed and applied in the system. BitCAT is also provided with a method to determine geometric constraints between the primitives during composition processes. BitCAT can make very useful CAD models for manipulation and NC machining, because they are properly attached with geometrical features. A demonstration verifies the faculty and usefulness of the system.1.

A demonstration verifies the faculty and usefulness of the system.1. IntroductionIn many areas of machine engineering, reverse engineering techniques can be used to create CAD models of real objects for which no CAD model exists. Reverse engineering typically starts with digitizing real objects so that surface or solid models can be deduced in order to exploit the advantages of CAD/CAM technologies. To digitize real objects, high energy X-ray computed tomography (CT) is very useful, because they can digitize whole shapes including interior regions accurately and automatically in a very short time [1].X-ray CT apparatuses obtain sectional images as two-dimensional (2D) bit-map data. After realobjects have been digitized by X-ray CT, they are represented as piled-up X-ray CT sectionalimages, that is three-dimensional (3D) bit-map data.

After realobjects have been digitized by X-ray CT, they are represented as piled-up X-ray CT sectionalimages, that is three-dimensional (3D) bit-map data. On the other hand, surface or solid models thatconventional CAD systems handle are represented as boundary representation (B-rep) data [2].Figure 1 shows the difference between B-rep and 3D bit-map formats. A 3D B-rep datum isrepresented as a set of geometrical and topological definitions of the shape elements like vertices,edges, faces, loops and shells. On the other hand, a 3D bit-map datum is represented as athree-dimensional array of tiny cells called “voxels” (volume picture cells). Each voxel holdsinside/outside discrimination or relative density at its position.

B-rep and bit-map formatsThe purpose of this study is to construct B-rep CAD models from 3D bit-map data obtainedfrom X-ray CT apparatuses. For this purpose, the authors developed a reverse engineering systemfor machine engineering, named “BitCAT”, which can make very useful B-rep models for CAD manipulation and NC machining. In this paper the authors describe the algorithm for creating B-rep CAD models and give results of an example conversion.2. Related workMany commercial software packages for reverse engineering create B-rep CAD models from 3D point cloud data obtained from 3D digitizers by the following procedures: (i) constructing triangulated meshes from 3D point cloud data; (ii) constructing a curve network, which divides a geometry defined by triangulated meshes into four-sided patches; and (iii) fitting NURBS surfaces onto each patch.Many mechanical parts include parametric primitives, such as flat planes, cylinders and swept surfaces.

Because the B-rep CAD models created according to the above approach are composed of NURBS surfaces only, parametric primitives in mechanical parts are represented as NURBSsurfaces. This causes the following problems.1. Because the data size required for handling a NURBS surface in CAD systems is muchlarger than that of a parametric primitive, CAD operations for such B-rep models mayneed a very long processing time.2. Making modifications for NURBS surfaces are much more difficult than for parametricprimitives.

Making modifications for NURBS surfaces are much more difficult than for parametricprimitives. For example, to change the diameter of a cylindrical hole represented asNURBS surfaces, the user must move many control points suitably.3.It is more difficult to generate NC tool paths for NURBS surfaces than for parametricprimitives.To overcome these problems, the authors suggest the following approach: (i) extracting240Voxels Voxelsregions from 3D bit-map data to which appropriate parametric primitives can be fitted; (ii) fittingparametric primitives to the extracted regions; and (iii) fitting NURBS surfaces to the regionswhere no appropriate parametric primitives can be fitted. Because parametric primitives arerepresented just as they are in the B-rep CAD models created by the authors’ approach, the B-repCAD models are very useful for CAD manipulation and NC machining.3.

Algorithm for creating B-Rep CAD modelsIn this study, the authors developed a reverse engineering system for machine engineering,BitCAT, which can construct B-rep CAD models from 3D bit-map data. The algorithm for creatingB-Rep CAD models is shown in Figure 2. First 3D bit-map surfaces are created from 3D bit-mapdata. Next regions are extracted from 3D bit-map surfaces to which appropriate parametric primitives can be fitted. Parametric primitives are fitted to the extracted regions. Then NURBS surfaces are applied on the regions where no appropriate parametric primitives are fitted. Then topological information of the B-Rep CAD models is created. Finally, the operator specifies geometric relationships among parametric primitives and refits them so that the specified geometric relationships can be satisfied. In some steps, interactive operations are required.

In some steps, interactive operations are required. Details of the algorithm in each step are described below.3.1 Creating 3D bit-map surfacesIn this stage, each voxel of 3D bit-map data is converted to a 1-bit voxel, which holds inside/outside discrimination, after thresholding has been applied. With a geometry represented as3D bit-map data 3D bit-map data1. Creating 3D bit-map surfaces 1. Creating 3D bit-map surfaces 2. Parametric primitives 2. Parametric primitives(Parametric primitives type, seed region) (Parametric primitives type, seed region)3. Fitting NURBS surfaces 3. Fitting NURBS surfaces (Curve network) (Curve network)4. Creating topological information 4. Creating topological information 5. Specifying geometric relationships and refitting surfaces 5.

Fitting NURBS surfaces (Curve network) (Curve network)4. Creating topological information 4. Creating topological information 5. Specifying geometric relationships and refitting surfaces 5. Specifying geometric relationships and refitting surfaces (Geometric relationships) (Geometric relationships)B-rep CAD models B-rep CAD models(............): user’s input (............): user’s inputFigure 2. Algorithm for creating B-Rep CAD modelsthe 3D bit-map data of the 1-bit voxel, the surfaces are composed of pixels on which an inside241voxel is adjacent to an outside voxel (Figure 3). Such pixels are called “surface pixels”. Surfacepixels are extracted from 3D bit-map data and 3D bit-map surfaces are created.3.2 Fitting parametric primitivesAfter 3D bit-map surfaces have been created, regions are extracted from 3D bit-map surfaces towhich appropriate parametric primitives can be fitted. Then parametric primitives are fitted to theextracted regions.

Then parametric primitives are fitted to theextracted regions. To accomplish fitting parametric primitives, the following three problems mustbe solved at the same time: segmentation of 3D bit-map surfaces, determination of parametricprimitive type and parametric primitive fitting [3].To solve these three problems, the authors developed a novel interactive “growing surfacemethod”. In the growing surface method, a user specifies the parametric primitive type and seedregion on 3D bit-map surfaces interactively using the BitCAT GUI interface. Then segmentation and fitting proceed automatically with the following procedures.First a surface of given parametric primitive type is fitted to the seed region. This surface which is called the “reference surface” is used to estimate how well a surface pixel matches the regions. Next the current region is growing by adding a surface pixel if the following requirements are satisfied.1.

Next the current region is growing by adding a surface pixel if the following requirements are satisfied.1. The surface pixel is adjacent to the current region. 2. The perpendicular distance between the surface pixel and the reference surface is within the threshold of D.3. The angle between the surface normal at the surface pixel and reference surface normal at the surface pixel position is within the threshold of S. The surface at the surface pixel is estimated by least-squares fitting of a neighboring ones.Outside voxel Outside voxelInside voxel Inside voxelThe region growing procedure is continueduntil no more surface pixels can be added to thecurrent region. Then the reference surface isrefitted to the current region. The above twoprocedures of region growing and surfaceSurface p ixel Surface p ixelrefitting are iterated mutually until the currentregion do not grow any further. Finally theFigure 3.

The above twoprocedures of region growing and surfaceSurface p ixel Surface p ixelrefitting are iterated mutually until the currentregion do not grow any further. Finally theFigure 3. Definition of surface pixelsparametric primitive is fitted to the current region.242In the authors’ current implementation, flat planes and cylinders are possible parametricprimitive types. To fit parametric primitives, a least-squares surface fitting is performed [4].3.3 Fitting NURBS surfacesIf there are regions where appropriate parametric primitives are not fitted, NURBS surfaces areapplied on the regions instead of them. This task is done by the following procedures: (i)constructing a curve network interactively using the BitCAT GUI interface, which divides regionsinto four-sided patches; and (ii) fitting NURBS surfaces on each patch [5].3.4 Creating topological informationIn this stage, topological information of B-Rep CAD models is created.

Example of creating topological information243region 4, region 5 and region 6 around the outer boundary and region 7 around the inner boundary.The edges can be defined by calculating intersecting curves between two adjacent regions. Thiscalculation is done according to information on the connection among the divided regions. In theabove example, edge 1 can be defined by calculating the intersecting curve between region 1 andregion 2 (Figure 4 (b) ).Then by linking the edges in the counterclockwise direction along the outer boundaries and inthe clockwise direction along the inner boundaries, the loops can be defined. As shown in figure 4(c), for example, loop 1 can be defined by linking edge 1, edge 2, edge 3, edge 4 and edge 5 in thecounterclockwise direction along the outer boundary.Next the vertices can be defined by calculating intersecting points between two linked edgesalong the loops.

For example, vertex 1 can be defined by calculating the intersecting point betweenedge 1 and edge 2, which are linked along loop 1 (Figure 4 (d)).3.5 Specifying geometric relationships and refitting surfacesIn many cases, mechanical parts have some geometric relationships among parametric primitives. For example, in the object shown in Figure 5, cylinder 1 is connected to plane 1 and plane 2 smoothly (C1 continuity). Plane 3 is perpendicular to cylinder 1, plane 1 and plane 2.Because each parametric primitive is fitted individually in the stage described in section 3.2, there are no geometric relationships among them yet. Then geometric relationships among parametric primitives are specified and the parametric primitives are refitted so that the specified geometric relationships can be satisfied.In the authors’ systems, a user specifies geometric relationships among parametric primitives interactively using the BitCAT GUI interface. Then parametric primitives are refitted.

Then parametric primitives are refitted. Geometric relationships among parametric primitives can be represented as constraint functions for the parameters defining the shapes. Forplane 3 plane 3example, the relationship that two flat planesare perpendicular to each other can berepresented as a constraint function for theflat planes’ normal vectors as follows:r ⋅ nn r 12=0plane 1 plane 1plane 2 plane 21nr and2nr are the normal vectors of flat planes. Then parametricwherethetwocylinder 1 cylinder 1Figure 5. Example of geometric relationships among parametric primitivesprimitives can be gotten which satisfy the244(a) (a)(b) (b)(c) (c)Figure 6. Example of BitCAT applied to 3D bit-map datumClose up Close up(b) (b)(a) (a)(b) (b)(c) (c)Figure 7. Example of BitCAT applied to 2D bit-map datumspecified geometric relationships by fitting them with some constraint functions.

Example of BitCAT applied to 2D bit-map datumspecified geometric relationships by fitting them with some constraint functions. Mathematically this task is equivalent to dealing with a constrained non-linear optimization problem. In this study, a quasi-Newton method [6] and Lagrange multiplier method [7] are used to solve this optimization problem.4. ExampleThe authors have applied the BitCAT system to some 2D / 3D bit-map data in order to confirmthe validity of the algorithms for creating B-Rep CAD models. The results are shown in thissection.First to test fitting parametric primitives and creating topological information procedures, the3D bit-map datum shown in Figure 6 (a) was used. This 3D bit-map datum was created by addingthe standard deviation of the noise to the synthetic data.

This 3D bit-map datum was created by addingthe standard deviation of the noise to the synthetic data. The data size was 100x100x100 voxels.Figure 6 (b) shows the complete B-Rep CAD models created by BitCAT and Figure 6 (c) shows theedges and vertices.245Next to test specifying geometric relationships and surfaces refitting procedures, the 2Dbit-map datum shown in Figure 7 (a) was used. This 2D bit-map datum was obtained from X-rayCT apparatuses. Figure 7 (b) shows parametric primitives (lines and circles) fitted to the 2Dbit-map datum, where no geometric relationships were specified. The authors specified C1continuity connections between parametric primitives. The results are shown in Figure 7 (c). Everyparametric primitive was connected with C1 continuity. This result is in 2D, but the algorithms canbe extended to 3D with no difficulty.5. ConclusionThe conclusions of this paper are summarized below;1.

This result is in 2D, but the algorithms canbe extended to 3D with no difficulty.5. ConclusionThe conclusions of this paper are summarized below;1. The authors have developed a reverse engineering system for machine engineering, named “BitCAT”, in order to construct B-rep CAD models from 3D bit-map data obtained from X-ray CT apparatuses.2. BitCAT can compose B-Rep CAD models by fitting parametric primitives, such as flat planes and cylinders. BitCAT is also provided with a method to determine geometric constraints between the primitives during composition processes. By using these functions, BitCAT can make very useful CAD models for manipulation and NC machining.3. BitCAT has been applied to some 2D / 3D bit-map data and the validity of the algorithms wasconfirmed.References[1] J. H Stanley and R. N. Yancey, CT-assisted Agile Manufacturing, Proceedings of SPIE,Nondestructive Evaluation for Process Control in Manufacturing (1996).[2] J. D. Foley, A. Dam, S. K.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDevelopment of Simulation Tools for Selective Laser Melting Additive ManufacturingThomas Nolan*, Yongsheng Lian*, Mark Sussman†Department of Mechanical Engineering, University of Louisville, Louisville, KY†Department of Mathematics, Florida State University, Tallahassee, FLAbstractTwo simulations tools have been developed to simulate selective laser melting. One is based on a multiphase flow solver with dynamic mesh adaptation and massive parallelism. The simulation tool is based on a first-principles approach to simulate complex additive manufacturing processes at the entire built part/component level. The developed model takes into account of heating, melting, powder-to-solid volumetric consolidation, cooling, as well as solidification and shrinkage that are often ignored in current simulation tools.

The second tool is to solve the heat equation only without considering the flow field and volumetric changes. In both tools laser is modeled as a heat source. The reported work is our first step toward the development of a complete software suite that can be executed rapidly on workstations and clusters with accelerators. The simulation tool can provide AM practitioners and researchers from industry and academic a fast and accurate simulation-based approach to replace current trial-and-error based practices in industry for process and material development.IntroductionLaser and metal powder based additive manufacturing (AM), specifically Laser sintering (LS) and Direct Digital Manufacturing (DDM), offers design freedom over the constraints of traditional manufacturing methods [1, 2]. LS is a process in which a high-energy laser beam scans a powder bed surface to melt the powder (ceramics, metal, or polymer) to form a bulk part [3].

LS is a process in which a high-energy laser beam scans a powder bed surface to melt the powder (ceramics, metal, or polymer) to form a bulk part [3]. The most common terminology for LS of metals is Selective Laser Melting (SLM) [4]. SLM can benefit from better design and lower cost, but LS is complicated due to the fast laser scanning rates and complex heat transfer and phase change processes [5-7]. There is a poor understanding of the important physical processes dealing with laser-material interactions, the heat transfer and molten metal flow, and metal phase transformations and thermal stresses, which hinders quality of a finished bulk part, such as density, dimensions, mechanical properties, and microstructure [7, 8]. Previous research also found the temperature field was inhomogeneous [9-13].

Previous research also found the temperature field was inhomogeneous [9-13]. And large thermal gradients increase residual stresses and deformation in metals and may lead to crack formation in a bulk part, which leads to the develop of serious problem in the SLM process [14].Due to the importance of temperature distribution in LS, researchers have grown towards understanding the SLM process [10, 15-19] and creating models to describe the thermal evolution of LS [5, 9, 11, 20-22]. Simulation models can demonstrate the influence of various parameters and are essential tools for identifying the proper parameters without expensive experimental testing.11220Various heat transfer processes occur in the SLM. Figure 1 shows the representation of heat transfer in SLM [23]. It consists of radiation at the powder bed, convection between the powder bed and the environment of an air void or gas filled chamber, and heat conduction inside the powder bed and between the bed and substrate.

The powder phase change is another important phenomenon in SLM, and phase change can significantly change thermal properties of the powder during SLM.Figure 1. Schematic representation of heat transferSLM metal temperature distribution has been modeled with heat equations without considering the flow field [13, 23-29]. Some researchers ignore the heat source as internal energy [30-32], but other studies put the source into the boundary condition equation [5, 20]. Other employ phase change and enthalpy to study the temperature field in SLM [5, 20, 30]. Another model considers influence from metal powder shrinkage and the molten pool fluid flow. Chan and Mazumder showed that fluid flow through the molten pool had a significant effect of the homogeneity of the sintering weld [33]. Some consider that the molten pool influences the thermal field in SLM. The model to simulate melting and solidification has been developed and used by [34-40].

Some consider that the molten pool influences the thermal field in SLM. The model to simulate melting and solidification has been developed and used by [34-40]. One- dimensional melting in mixed powders has been solved analytically by [37]. A three-dimensional model has been created that considers the thermal behavior and fluid dynamics in the molten pool caused by Marangoni and buoyancy forces along with the melt pool flow caused by gravity and capillary forces [39]. The model also accounts for powder shrinkage due to density and uses a fixed grid temperature transfer model for modeling the melting and solidification of the powder. The model has been compared to the cross-sectional profile for the melt-solid interface in laser processing for nonporous metal experiments.

The model has been compared to the cross-sectional profile for the melt-solid interface in laser processing for nonporous metal experiments. The results show that the thermocapillary force, shrinkage, and buoyancy force dominate the melting and fluid flow [39].Another thermal evolution of SLM as a heat transfer process utilizes the Fourier heat conduction theory with a Gaussian laser source embedded as an internal energy and the volume-of-fluid method to study the multiphase interaction forces and temperature variation in Ti-6Al-4V powder. Unlike the model from [39], the Marangoni and buoyancy forces are not considered. The latent21221heat is dependent on temperature and considers liquid phase of the powder particles. The model uses a volume fraction transport equation to present the distribution of metal and gas phase within the computation domain.

The model uses a volume fraction transport equation to present the distribution of metal and gas phase within the computation domain. The model has been compared to a cross-sectional profile for the distribution of the melting pool.Various models have been proposed to model the heat transfer from laser beam. Some models have sought to use the volumetric laser beam distribution assumption. Niebling and Otto considered a volumetric line heat source [41]. Gusarov and Yadroitsev used a radiation transfer equation with an isotropic scattering term to describe laser beam penetration [30]. Although the laser beam penetration problem has been studied, Kolossov and Boillat found that with an element size larger than five gradient diameters, the laser beam penetration can be ignored [12]. Realizing the interaction between the laser beam and powder bed, leads to a greater understanding of laser beam penetration and absorption.

Realizing the interaction between the laser beam and powder bed, leads to a greater understanding of laser beam penetration and absorption. Several models consider the absorption a constant ratio for the bulk powder [20, 42-46]. Laser beam absorption in powder depends on several factors including oxidation level, surface type, surface temperature, and laser beam wavelength. In the case of metal powders, the laser beam absorption ratio varies within a few percent of the molten absorption ratio [47].Thermal properties of metal powder include density, enthalpy, heat capacity, and thermal conductivity. Carslaw and Jaeger show that thermal conductivity varies with temperature [48]. SLM uses an effective thermal conductivity. It has also been shown that the porosity of the powder bed influences thermal conductivity.

SLM uses an effective thermal conductivity. It has also been shown that the porosity of the powder bed influences thermal conductivity. The effective thermal conductivity is a function of the thermal conductivities of solid metal and gas in the powder voids [3, 20, 49].The main consideration of thermal conductivity of metal powders is the temperature near the melting point and above in liquid state. Without experimental data, a constant thermal conductivity is assumed. Thummler proposed that thermal conductivity is influenced by powder porosity and pore geometries and thermal conductivity is controlled by gas content in the voids of the powder [50]. Rombouts et al. showed that thermal conductivity is independent of powder material and depends on porosity and size of voids in the powder along with depending on gas content in the voids [51].

Rombouts et al. showed that thermal conductivity is independent of powder material and depends on porosity and size of voids in the powder along with depending on gas content in the voids [51]. Roberts and Wang show that the thermal conductivity of Ti6Al4V powder is low for powder and increases dramatically when nearing the melting point [20]. The density and heat capacity of metal powder is shown as the mass average of each state’s (powder, liquid, solid) value [49]. Some consider the heat capacity to be separate at each phase and use an effective heat capacity that is temperature dependent [46].In this paper, we present two models to study SLM. One is based on commercial software Fluent, and the other one is based on an in house multiphase flow solver. In the first approach, we solve the heat transfer equations without considering the flow field. The laser is modeled as a three- dimensional Gaussian energy source that is imposed on a two-dimensional surface.

The laser is modeled as a three- dimensional Gaussian energy source that is imposed on a two-dimensional surface. This approach only considers the heat transfer without the phase change. In the second approach, we will model the phase change using multiphase flow solver.Numerical Models31222In the following we introduce the two models used to study SLM. The first one is based on a simple heat equation (the heat conduction model) and the second one is based on a multiphase flow solver (multiphase flow model).Heat Conduction ModelThe most common thermal evolution of SLM as a heat transfer process can be described by Fourier heat conduction theory.

Equation (1) describes the governing heat conduction in a moving medium [48],𝜆 (𝜕2𝑇 𝜕𝑥2 +𝜕2𝑇 𝜕𝑦2 +𝜕2𝑇 𝜕𝑧2 ) + 𝑞 = 𝜌𝑐𝜕𝑇 𝜕𝑡The initial condition and boundary conditions are as follows:Initial temperature:𝑇(𝑥, 𝑦, 𝑧, 0) = 𝑇0Surface convection and radiation:−𝜆𝜕𝑇 𝜕𝑧= 𝜀𝜃𝜎(𝑇4 − 𝑇𝑒4) + ℎ(𝑇 − 𝑇𝑒)No heat loss at the bottom:−𝜆𝜕𝑇 𝜕𝑧|𝑧=0 = 0where 𝑇 is the temperature, 𝜆 the conductivity coefficient, 𝜌 is the density, 𝑐 is the heat capacity coefficient, 𝑞 is the internal heat, 𝑇0 is the powder bed initial temperature, 𝑇𝑒 the environment temperature, 𝜀𝜃 the thermal radiation coefficient, 𝜎 the Stefan-Boltzmann constant, and ℎ the convection heat transfer coefficient. The governing equations were used to study the thermal field of dental porcelain in SLM [47].Multiphase Flow ModelThe second approach is based on a multiphase flow solver.

In this approach the governing equations to be solved are the three-dimensional Navier-Stokes equations for immiscible, multiphase flows, which are given as follows:0VTtensionVVVpVVFgtvTcVTqtwhere,,Vuvwis the velocity vector, ρ is the density, t is the time, p is the pressure, µ is theviscosity,gis the gravitational acceleration vector,qis the heat flux vector.Besides the above temperature condition at the interface, the following mass conservation must be satisfied across the interface [52]41223(1)(2)(3)(4)(5)(6)(7)LILsIsmVnVnVnVn(8)wheremis the interfacial mass flux,Landsare the densities of liquid and solid phasesVnrespectively, and satisfies the Rankine-Hugoniot jump conditionis the normal velocity of the fluid, is the normal velocity of the interface.

The energy conservation at the interface equationnis the normal pointing from solid to liquid,IVn()LsLsmhhqnqn(9)whereLhandshare the enthalpy of the liquid and solid phases respectively.The liquid-solid interface can be represented by a zero level set of a distance function. The distance equation is given by the following equation [52]0 or 0LLmDmVtDt(10)where is the distance to the piecewise linear moment of fluid reconstructed interface. The multiphase flow solver solves the three-dimensional Navier-Stokes equations using the variable density pressure projection algorithm [53] on block structured adaptive mesh refinement grid (AMR) [54]. It can handle both compressible [55] and incompressible flows [56, 57]. The solver employs the state of the art moment of fluid method (MOF) [58-60] to represent multiphase interfaces [55, 61-64]. It employs dynamic contact models for droplet impact problems [65].

The solver employs the state of the art moment of fluid method (MOF) [58-60] to represent multiphase interfaces [55, 61-64]. It employs dynamic contact models for droplet impact problems [65]. Our tests show that the code has a high parallel efficiency of more than 96% on a 48-core workstation.Interface representationDuring the MOF interface reconstruction process, a reference volume fraction functionrefFand areference centroid,,crefxboth corresponding to the real interface, are given [Figure 2a], and theactual volume fraction function, FA, and the actual centroid, interface, are then computed [Figure 2b].,cAxcorresponding to the reconstructedabFigure 2. Surface reconstruction using the MOF method [56]. (a) the real interface and reference volume fraction and centroid; (b) the reconstructed interface and computed volume fraction and51224centroid.

(a) the real interface and reference volume fraction and centroid; (b) the reconstructed interface and computed volume fraction and51224centroid. The MOF reconstruction requires the reconstructed volume to be equal to the reference volume while the reconstructed centroid as close to the reference centroid as possible.The MOF method requires that the actual volume fraction be equal to the reference value and the actual centroid be as close to the reference centroid as possible.

This becomes a constraint optimization problem formulated as follows:2Minimize , under constrain: ,0ccrefArefAbFFbxxnnAs discussed in [57], the use of centroid information ensures the MOF method maintains a sharp interface which is critical for the droplet impact simulation because the thickness of the lamella is about two orders of magnitude smaller than the initial droplet diameter [66].In our numerical models, we used two different materials to differentiate the liquid drop and the liquid film to illustrate the pertinent fluid dynamics features in the drop impact on thin film. The method is illustrated in Figure 3. We first reconstruct the interface for the material whose centroid is furthest the centroid of the computational cell by solving Eq. (13). In the case shown in Figure 3 it is the red material.

We first reconstruct the interface for the material whose centroid is furthest the centroid of the computational cell by solving Eq. (13). In the case shown in Figure 3 it is the red material. Next, we reconstruct the interface for the remaining material whose centroid is furthest to the centroid of the unoccupied region in the cell (white region) by solving Eq. (13). In Figure 3 the next material is marked in blue. This procedure continues until all the material interfaces are constructed.Figure 3. MOF interface reconstruction with three materials in one cellAdaptive mesh refinementThe Navier-Stokes equations are solved using the variable density pressure projection algorithm [53] on the block structured adaptive mesh refinement grids [54, 67]. The grid adaption is based on the triple point region and the curvature of the interface. As shown in Figure 4, the grid refinement is performed near regions where curvature is greater than a predefined value.

As shown in Figure 4, the grid refinement is performed near regions where curvature is greater than a predefined value. The adaptive mesh refinement method ensures fine grid is only used in the regions of interests, which maintains the accuracy of the solver at reasonable computational cost.61225(11)Figure 4. Two-level adaptive mesh refinementLaser Source Term ModelThe modeling of laser beam characteristics has been researched to better understand the LS process. Simple models assume the laser beam to be a point source, but this conjecture is not representative of reality. It has been shown that laser beam sources can be characterized by three parameters: diameter, power, and intensity distribution. Courtney and Steen deduced an effective Gaussian beam diameter equivalent using measurements from a laser beam and comparing to a Gaussian laser beam diameter [68]. This effective Gaussian laser beam diameter is the most widely used for simulation [20].

This effective Gaussian laser beam diameter is the most widely used for simulation [20]. In this model the laser intensity distribution can be written as221(/)0()ddIrIewhere d1 is the beam diameter, I0 is the irradiance that diminishes exponentially, and d is the radius of a point from the center of the laser beam source. The thermal heat flux distribution is𝑞(𝑟) =2𝑃 𝜋𝑟02 𝑒−2𝑟2 2 𝑟0where P is the laser power, 𝑟0 is the spot radius, and 𝑟 is the radial distance. The total heat flux on a circular surface of radius r0 can be calculated by integrating the heat flux𝑞𝑚 =𝛼 𝜋𝑟0𝑟02 ∫ 𝑞(2𝜋𝑟)𝑑𝑟 =00.865𝛼𝑃 𝜋𝑟02where 𝛼 is absorption rate.The laser beam distribution is typically assumed to be either surface or volumetric. The surface variation of the laser beam is the most common because of the lack of research into laser beam penetration in the volumetric distribution [9, 30, 69].

The surface variation of the laser beam is the most common because of the lack of research into laser beam penetration in the volumetric distribution [9, 30, 69]. In an attempt to solve the lack of understanding of the volumetric distribution, Wang and Laoui used a ray tracing (RC) model that predicts the absorption and reflection of laser beam energy on each particle produced by a large71226(12)(13)(14)number of rays [9]. The energy absorbed by each particle is totaled and the total energy delivered by the laser is totaled. From the total energy calculation energy, penetration can be described as a depth function into the powder.Li at el. used the assumption that the laser beam distribution is surface in nature, but also includes the convection and radiation terms from the Fourier heat conduction theory [46].

used the assumption that the laser beam distribution is surface in nature, but also includes the convection and radiation terms from the Fourier heat conduction theory [46]. The laser source model is𝑆𝑙𝑎𝑠𝑒𝑟(𝑥) = [𝛼𝑃 𝜋𝑟02 exp [−2(𝑥 − 𝑣𝑡 − 𝑥𝑖)2 2 𝑟0] − 𝜀𝜃𝜎(𝑇4 − 𝑇𝑒4) + ℎ(𝑇 − 𝑇𝑒)] 𝛿where 𝛼 is the absorptivity, 𝑃 is laser power, 𝑟0is the laser beam radius, 𝑣 is scanning velocity, 𝑥 is the laser beam position, 𝑥𝑖 is the initial laser beam position, 𝜀𝜃 is the black body radiation, 𝜎 is 4 is environment temperature, and 𝛿 is a delta Boltzmann’s constant, 𝑇 is melt pool temperature, 𝑇𝑒 function. The purpose of the laser source equation is to mark the position of the powder particle surface in the flow field. The laser source is implemented at the beginning of simulation until end of the laser beam path.Numerical ResultsThe model represents IN718 powder on top of a steel plate with air surroundings. The process parameters and thermal properties are taken from [46] and [70].

The process parameters and thermal properties are taken from [46] and [70]. Table 1 shows the model and process parameters. ANSYS Fluent is used to create the simulation. The laser beam is modeled as Gaussian distribution with a 100 μm laser spot size. The element size is 100 μm in a 0.8 cm by 0.4 cm by 0.226 cm computation domain and 73600 elements. The laser scans from left to right. Convection and radiation are modeled using a mixed heat transfer coefficient at the powder surface included in the source term. The initial temperature is set at 300 K. The density and thermal conductivity are functions of temperature. The specific heat is constant.The model is set up using a rectangular solid Ti6Al4V part with mesh and a user defined laser beam source beam function taken from [46]. The laser source is implemented at the start of the simulation and continues until the laser beam moves over the entire path of the domain.The mesh is shown in Figure 5.

The laser source is implemented at the start of the simulation and continues until the laser beam moves over the entire path of the domain.The mesh is shown in Figure 5. The temperature profile for the first layer scan is shown in Figure 6., and Figure 7.. The temperature distribution contours show an ellipse-like shape that repeats as the profile moves from the center, which is shown in [71]. The melt pool moves linearly maintaining the ellipsoid shape throughout. The melt pool also maintains a similar shape and temperature concentration at each time instant. The plots also show the concentration of the thermal gradient at the front of the molten pool, which is due the movement of the laser beam source and the temperature and phase dependent material properties.Figure 8 shows a comparison of the experimental results to our simulation. The center and shape of the temperature field in our simulation are like those of the experimental results.

Model and process parametersIn Figure 9 we show the simulated melting process using our proposed multiphase flow solver. Three metal particle of the same radii of 30 𝜇𝑚 are put side by side. The laser source model of Eq. (15) is used. The laser power is 175 W and scan speed is 1.25 m/s. The density and thermal conductivity of the Ti6A14V are assumed to be constant. The melting process is captured by the solver. Our simulation also shows that the particle is not fully melted.101229Figure 9. Simulated melting process based on the multiphase flow solver. (Red: Solid Ti6A14V).ConclusionsWe presented two models to study SLM. The first model is based on the simple heat conduction equation that considers the heat transfer but not the flow field or volumetric changes.

The first model is based on the simple heat conduction equation that considers the heat transfer but not the flow field or volumetric changes. The other model is based on a multiphase flow solver that considers the heat transfer, phase change and volumetric changes.AcknowledgementThis work is partially supported by a NASA Grant.1112301.2.3.4.5.6.7.8.9.10.11.12.13.14.15.16.17. 18.19.20.ReferencesZhong, M. and W. Liu, Laser surface cladding: The state of the art and challenges. Journal of Mechanical Engineering Science, 2010. 224(C5): p. 1041-1060. Zhou, J.H., Y.W. Zhang, and J.K. Chen, Numerical simulation of random packing of spherical particles for powder-based additive manufacturing. Journal of Manufacturing Science and Engineering: ASME, 2009. 131(3). Cervera, G.B.M., Numerical prediction of temperature and density distributions in selective laser sintering processes. Rapid Prototyping Journal, 1999. 5(1): p. pp. Gibson, L., D.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDevelopment of Soft Gripper Pneumatic Control System Based on Deep Reinforcement LearningSeongyeon Kim, Kiseong Kim, Jongho Shin*Department of Mechanical Engineering, Chungbuk National University, Cheongju, Republic of Korea 28644AbstractAs interest in soft grippers soared, many studies have been performed to control the soft gripper. For the soft gripper control, a soft gripper model is required first. Usually, the soft gripper modeling has been done through finite element analysis, which takes lots of time and is effective only in limited situations. Therefore, research on deep learning-based modeling with a small amount of FEM results has been extensively conducted, and some satisfactory results have been reported.

Therefore, research on deep learning-based modeling with a small amount of FEM results has been extensively conducted, and some satisfactory results have been reported. However, since the model is expressed in the form of a neural network, it is difficult to utilize general control methods, so research on optimal control or deep reinforcement learning is being attempted. In this study, we propose a pneumatic control system for the soft gripper control based on the DRL. To this end, the soft gripper and DRL-based controller are directly developed, and experiments are performed and the results are analyzed.IntroductionAs interest in a soft gripper has recently increased, many studies have been conducted to control the soft gripper[1]. The soft gripper model is required for soft gripper control. In general, soft gripper modeling is done through nonlinear finite element analysis. However, this takes a lot of time and is effective only in limited circumstances.

In general, soft gripper modeling is done through nonlinear finite element analysis. However, this takes a lot of time and is effective only in limited circumstances. Active research on deep learning-based soft gripper modeling using a small amount of nonlinear finite element analysis results has been conducted. However, the deep learning model is expressed in the form of a neural network, making it difficult to use general control methods[2]. Research on optimal control and control through deep reinforcement learning has been attempted. In this study, we propose a pneumatic control system for soft gripper control based on deep reinforcement learning.Fig 1. The soft gripperThe pneumatic control systemAs shown in Fig 2, The proposed pneumatic control system consists of hardware and software. First of all, the hardware was equipped with the soft gripper so that it could be driven.

First of all, the hardware was equipped with the soft gripper so that it could be driven. The overall hardware configuration is shown in Fig 2, and it was possible to move the axis up and down, left and right to create various actuating and measurement environments. The pneumatic system for actuating the soft gripper was designed and manufactured. A small compressor was used as the pneumatic source, and the available maximum air pressure was 6 bar. A filter module was installed to remove moisture and foreign substances from the compressed air generated from the compressor.1052Fig 2 Conceptual diagram of the soft gripper pneumatic control systemThe motorized regulator used in this study is a ITV-1090 model of SMC, and was controlled by a PC. This system generates operation commands through software composed of GUI and controller, and the entire communication is composed of RS232C.

This system generates operation commands through software composed of GUI and controller, and the entire communication is composed of RS232C. In addition, the load applied to the load cell can be monitored in real time through the operation of the soft gripper. since the interface is already configured, soft gripper control is sufficiently possible simply by changing the controller part to a DRL-based controller.Fig 3. The soft gripper pneumatic control systemOn the other hand, the software was configured using the VB.net language in Microsoft’s Visual Studio environment. As shown in Fig 4, The graphical user interface (GUI) was created for user convenience. The user can set a target pressure, the time taken to reach the target pressure, and the number of repetitions of the experiment can be automated through these settings.Fig 4.

The user can set a target pressure, the time taken to reach the target pressure, and the number of repetitions of the experiment can be automated through these settings.Fig 4. The GUI for soft gripper pneumatic control system1053In addition, the air pressure currently supplied to the soft gripper and the pressure applied to the load cell by the gripper can be monitored in real time.The experiment resultsFig 5. Correlation between a supply pressure and the load applied to the load cellAs shown in Fig. 5, the soft gripper actuating experiment was performed using the constructed system, and it was confirmed that this system operated normally. The proposed system is easy to access by controlling the entire system through GUI. Therefore, various functions of the soft gripper can be verified and tested smoothly.

The proposed system is easy to access by controlling the entire system through GUI. Therefore, various functions of the soft gripper can be verified and tested smoothly. Furthermore, This system is evaluated to be effective in actuating and controlling various types of soft grippers as an integrated form of hardware and software.ConclusionThe proposed system was built to utilize the DRL-based controller when controlling the soft gripper, and the possibility was confirmed through driving experiments. Therefore, it can be effectively used for actuating and controlling the soft gripper through the DRL-based controller in the future.AcknowledgementThis work was supported by Basic Science Research Program through the National Research Foundation of Korea(NRF) funded by the Ministry of Education (No.2021R1A4A1033141)References[1] Kim, Daekyum, et al. "Review of machine learning methods in soft robotics." Plos one 16.2(2021): e0246102.

VELOPMENT OF TECHNIQUE FOR 3D PRINTED MOULD INTRICATE RAPID CASTINGDermot Brabazon, Donal Kennedy and Michael TyrellSchool of Mechanical and Manufacturing Engineering, Dublin City University, IrelandAbstractThe development of rapid prototyping (RP) technologies has proven highly significant in the efforts to reduce production times for a number of casting processes. Although rapid prototyping technologies have most commonly be used to produce patterns for investment and sand casting this study demonstrated the use of direct shell production casting using three dimensional printing technology. Statistical methods as well as thermal, visual and dimensional analysis were also applied in order to optimise this rapid casting (RC) process for thin wall non-ferrous parts. Mould dimensions within ± 0.22 mm were achieved with the developed technique. Higher melt temperatures and pouring pressures produced castings with better dimensional accuracy.

Mould dimensions within ± 0.22 mm were achieved with the developed technique. Higher melt temperatures and pouring pressures produced castings with better dimensional accuracy. Mould temperature was not found to affect the casting dimensional accuracy significantly. The results allow for better dimensional specification of CAD file geometry for the rapid casting process.IntroductionIn the last thirty years great strides have been made in developing and adapting processes for higher production rates and reduced lead times. Many useful rapid prototyping (RP) technologies are now available for product development [1]. The advent of rapid prototyping technologies such as selective laser sintering and 3D printing has contributed to this significantly, particularly in the area of non-ferrous precision casting. The use of these new technologies in this respect has been titled rapid casting.

The use of these new technologies in this respect has been titled rapid casting. An example of the use of this technology is in the production of solid metal moulds, termed Rapid Tooling (RT). The process under examination in this study was direct shell production casting using a 3D printing machine, a technology originally developed by MIT. A number of companies now supply variations along the lines of this original 3D printing technology. In a similar manner to investment casting, but without the wax and sintering, the mould can be printed on the 3-D printer machine and then cast into to produce the final part. New sand based material compositions have also been developed for rapid casting with these machines.These methods should allow the potential to produce thin wall sections while maintaining high dimensional accuracy.

Shrinkage of the rapid prototyped component leading to dimensional inaccuracies is a problem with many RP techniques [2] and gas inclusion is a common problem with conventional plaster moulding. Hygroscopic expansion of the investment and the thermal contraction of the cast metal can lead to dimensional changes [3]. Even on heating or cooling of a mould thermo-mechanical effects can cause dimension changes. Previous workers have presented finite element assessments of the affects of thermal dimensional changes on the tolerance of ceramic mould usage for RT [4]. Appropriate selection and control of the printing process, metal pouring temperature and proper design of the gating system can largely overcome these problems [5-10]. Recent work has shown possibility of using 3D printing for successful large scale production of aluminum, lead and brass alloys [11-13].

Recent work has shown possibility of using 3D printing for successful large scale production of aluminum, lead and brass alloys [11-13]. Despite these previous studies, there is a lack of work presented in the literature on the possibilities of achieving dimensionally accurate thin sectioned castings using 3D printed plaster moulds. The primary objective of this study wasto determine optimal parameters for the 3D printing of plaster moulds and the associated casting process to ensure correct mould filling of thin sectioned moulds.Materials and methodsThe alloys used were aluminium alloy A356 and Sn-10%Pb solder. Initial experiments proceeded using the aluminium alloy. The tin-lead solder was then used in a statistical design of experiments which although a far less industrially useful metallic alloy with respect to casting applications proved useful in inferring the results onto other alloy systems.

The critical process parameters selected for examination were filling pressure, mould preheat temperature and melt superheat temperature. The design of experiments method was used to arrive at the more optimal values for the process. A two level three-parameter factorial analysis (23) was carried out resulting in eight experiments in total. Response variables chosen for analysis were measured dimensions of the produced castings. A Vernier callipers and micrometer screw (both accurate to +/-0.01mm) were used for all measurements. The casting dimensions were related to the mould dimensions and CAD file dimensions. The effects of each factor on the response variables were analysed. Mould design and fabricationA four-stepped cavity mould design which was used in this work is shown in Figure 1. The first step section, nearest the pouring basin, was 8mm thick. Each subsequent section was 2 mm less thick (providing 6, 4, and 2 mm thick steps respectively).

The first step section, nearest the pouring basin, was 8mm thick. Each subsequent section was 2 mm less thick (providing 6, 4, and 2 mm thick steps respectively). Each step was 35 mm long and 100 mm wide. The total cavity length was 140 mm. Wall thickness of the surrounding cavity (including a printed flat plaster base plate) was nominally set at 10 mm. The wall thickness was larger by the amount indicated above for the smaller steps. A 40 mm diameter sprue was used. Three millimeter diameter holes, central to each step, which ended four millimeters away from the cavity, were incorporated in the mould. These were used to take temperature measurements along the length of the mould during pouring. Type K thermocouples were used for these measurements. Two vent holes, one millimeter in diameter, were also incorporated at either side of the last step in the mould cavity.

Type K thermocouples were used for these measurements. Two vent holes, one millimeter in diameter, were also incorporated at either side of the last step in the mould cavity. Figure 2 shows a picture of the mould, thermocouples and casting set-up.The machine used to manufacture the plaster moulds was the Z402 three-dimensional printer. The 3D printer used STL files of the CAD mould design to produce the plaster parts. The mould production time was typically about 3.5 hours. Once the machine cycle was complete the part was left to dry for between 30 minutes to 1 hour before being depowdered. The moulds were left to air-dry for one day to obtain sufficient strength before casting. MaterialsThe principal mould ingredient was calcium sulphate (plaster) powder. Calcium sulphate consists of two different hydrate forms and an anhydrous form.

MaterialsThe principal mould ingredient was calcium sulphate (plaster) powder. Calcium sulphate consists of two different hydrate forms and an anhydrous form. The dihydrate form (CaSO4-H2O) or gypsum when heated above 128 °C (equilibrium temperature) loses 75% of its water to form the hemi hydrate (CaSO4-0.5H2O), plaster of paris. When it is heated above its equilibrium temperature of 163 °C all of the water of crystallisation is removed producing the anhydrous calcium sulphate (CaSO4). The 3D printing machine printed binder to bind the powder in consecutive 2D layers to build up the mould. The alloys used for these tests were solder alloy Sn- 10%Pb and aluminium alloy A356 (Al-7%Si-0.4%Mg). The high silicon content in A356 provides for excellent casting properties including fluidity.

The alloys used for these tests were solder alloy Sn- 10%Pb and aluminium alloy A356 (Al-7%Si-0.4%Mg). The high silicon content in A356 provides for excellent casting properties including fluidity. The melting range (solidus to liquidus temperature) of the Sn-10%Pb alloy was 183 to 235 °C and that of A356 was 572 to 614 °C.Figure 1 Mould design used in casting experiments.Figure 2 Mould, thermocouples and casting set up.Casting conditionsTwo furnaces were used in the casting process. One was used to dry and preheat the moulds while the other was used to melt the alloy. The moulds were heated for one and a half hours at the preheat temperature to ensure that all moisture had evaporated and that the mould temperature was homogenous. For the casting trials, a 3 kg weight was placed on top of the mould and fire clay was used to seal the mould edges. In the aluminium casting trials, a casting temperature of 725 °C was found to result in successful castings being formed.

In the aluminium casting trials, a casting temperature of 725 °C was found to result in successful castings being formed. Initial tests determined that the minimum temperature at which the Sn-Pb melt could be successfully pouredwithout premature freezing was 235 °C. The time for which the alloy was heated was also approximately one and a half hours, half an hour to reach temperature and one hour to ensure temperature homogeneity. Pouring time was approximately 8 - 10 seconds. Surface dross was skimmed off before casting.Table 1 shows the set of parameters examined and the levels at which each parameter was tested for the Sn-Pb design of experiment castings. The response variables used were the dimensions of the mould and casting such as the length to which the metal filled into the last step. This was quantified by measuring the average length that the melt flow progressed during casting.

This was quantified by measuring the average length that the melt flow progressed during casting. This average was calculated as the average of five distance measurements taken along the casting (side lengths, middle length and length measurements half way between these). The locations at which other measurements were taken are indicated in Figure 3. Measurements of step heights were also taken and are denoted herein as D1, D2, D3 and D4, for step heights 1 to 4 respectively.

Measurements of step heights were also taken and are denoted herein as D1, D2, D3 and D4, for step heights 1 to 4 respectively. The measurement process involved taking three repeated measurements at these positions in ten repeated mould fabrications and corresponding positions in the cast part.Table 1 Levels of factors used for design of experimentsA: Mould Temperature (°C) B: Melt Temperature (°C) C: Pressure (Pa) *Low Level 158 240 1379High Level 183 250 5517Low pressure: 20 mm melt head and high pressure: 80 mm melt headFigure 3 Plan view of mould cross section with reference dimensional tags.Results and DiscussionDuring initial aluminium casting trials, some tarnishing of the cast surface resulted due to burning of the mould in contact with the metal. After testing at different temperatures in the furnace it was noted that, if left for sufficiently long periods, burning on the surface of the mould was seen to occur at temperatures above 200 °C.

After testing at different temperatures in the furnace it was noted that, if left for sufficiently long periods, burning on the surface of the mould was seen to occur at temperatures above 200 °C. This tarnishing of the casting was overcome by applying very thin (< 0.1 mm thick) surface coatings of graphite in one instance and zircon in another to the mould surface. Although the mould colour faded during casting, it did not excessively burn and the mould coatings protected the casting surface. It was also clear from these aluminium casting tests that the positioning of the vents at the end of the mould was critical to mould filling in order to allow evolved gases to escape. Two one millimeter through vents were placed at the end corners of the mould for all subsequent tests. The lower temperature Sn- 10%Pb alloy was also used for the rest of the tests presented below.

Two one millimeter through vents were placed at the end corners of the mould for all subsequent tests. The lower temperature Sn- 10%Pb alloy was also used for the rest of the tests presented below. Mould burning did not occur excessively in this case due to the lower pouring temperatures and so no mould wash (thermal barrier) was used. Visual InspectionAs occurs with conventional plaster casting, the mould material adhered to the surface of the cast part. After the cast part cooled, water was used to convert the hemi-hydrate plaster material to its soft hydrate form which was easily detached from the cast part. Dross appeared in the gating system and the castings were judged to be largely free of surface dross. A picture of one of the castings is shown in Figure 4.

Dross appeared in the gating system and the castings were judged to be largely free of surface dross. A picture of one of the castings is shown in Figure 4. The castings picked up the mould features very well as is typical of plaster casting.Figure 4 Isometric picture of casting run number three which was processed at the low mold temperature, high pouting temperature, and with low pressure.Thermal measurements Temperature profile recordings confirmed the amount of time required for each casting was between within 8 to 10 seconds. A thermocouple which was pierced though the mould also indicated a steady 70 ºC temperature difference between the solidifying metal in the mould cavity and the thermocouple four millimeters away in the plaster mould.Dimensional Accuracy The dimensions of produced moulds and castings were compared to the corresponding CAD file dimensions.

Table 2 shows the average amount by which the mould and casting dimensions differed from those specified above in the casting conditions section above. The measurements taken from the moulds reveal that the average dimensional accuracy of the printing process was ± 0.22mm. The thin sections of the mould were found to be smaller than corresponding CAD file dimensions. For example, the mould walls were approximately 0.22 mm less than specified in the CAD file. The increase in the mould cavity internal dimensions could also be attributed to this mould contraction and warping. Where the mould filled fully, the casting dimensions were slightly larger than the corresponding CAD file dimensions. The widths and lengths of steps were in general of similar measurement and accuracy to those found in the fabricated moulds. The thickness of each step was however approximately 30% greater in the castings compared to the CAD file dimensions.

The thickness of each step was however approximately 30% greater in the castings compared to the CAD file dimensions. The length of fill of the last step was dependant on the casting conditions. The dimensions of the last step are shown in Table 3 for each of the design of experiments casting conditions. The final step (step 1) fill distance (L1) was found to be the best discriminator between different casting conditions. The final step fill distance, width and thickness were lower at the lower pressure and superheat levels.

The software used to analyse the data was Dataplot, a data analysis program. The coefficients for the above equation are shown in Table 4 and the degree of fit is shown in Table 5. The sum of the square total, equation 2, and that fraction of the variance, equation 3, associated with the each factor demonstrates the level to which the factor effects the response. The F- statistic is the mean square for the factor divided by the mean square for the error.Table 4 Coefficient effect table where T = q/σ (effect/standard deviation).Identifier Mean C B BC AC AB A ABCMean Effect (q coefficient) 31.06500 6.14500 4.23 -3.255 -2.145 1.97 1.17 -0.745T Value 8.2 5.7 -4.4 -2.9 2.6 2.6 1.6 -1.0The results from statistical analysis show that the most significant factor in effecting the mould fill level was the pressure followed by the superheat level. This was apparent from the dimensional measurement analysis of each casting as presented in casting dimension results section above.

This was apparent from the dimensional measurement analysis of each casting as presented in casting dimension results section above. From the residual sum of squares and F statistic values (Table 5) for each factor itis also clear that random error in the experiment had a significant effect in the experiment. The reliability of these results must therefore not be overestimated.ConclusionsMould dimensions were found to be accurate to within ± 0.22 mm for all the moulds produced. This value is higher than the best to which the 3D printing process is capable of, now less than 0.1 mm. The accuracy that can be obtained depends however on the mould dimensions. Combinations of thick and thin sections can lead to loss of dimensional accuracy due to shrinkage and warpage. The usefulness of the 3D printing process in developing plaster moulds for rapid casting has been demonstrated.

The usefulness of the 3D printing process in developing plaster moulds for rapid casting has been demonstrated. A model has been generated which allowed the relation between processing parameters and casting success to be evaluated. This was measured by the extent to which the process produced casting dimensions close to those of the original CAD file dimensions. Statistical analysis of the results revealed that the dominating factor affecting the extent of mould fill in the casting process was the pouring pressure. The next most significant factor was the melt superheat. The mould temperature was not found to have much effect on casting dimensions. However, the strength of the mould material was seen to vary with temperature which may impact on casting dimensions for higher pressure castings. Further work is also necessary to extend the range of materials to which this process can be applied.References 1. D.T. Pham and R.S.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDevelopment of virtual surgical planning models and a patient specific surgical resection guide for treatment of a distal radius osteosarcoma using medical 3D modelling and additive manufacturing processesMazher Iqbal Mohammed*, Mark Gerard Ridgway and Ian GibsonDeakin University, School of Engineering, 75 Pigdons Road, Waurn Ponds, Geelong, VIC 3216, AustraliaKeywords: 3D printing, Resection Guide, Radius, 3D modelling, ImplantCorresponding Author: mazher.mohammed@Deakin.edu.auAbstractIn this study we will assess the design and fabrication of a patient-specific resection guide to augment surgical procedures, such as bone grafts and implant placement. Medical imaging data was used to form a 3-dimensional, digital template model of the target anatomy to incorporate surface topography information into the guide.

Medical imaging data was used to form a 3-dimensional, digital template model of the target anatomy to incorporate surface topography information into the guide. The surgical guide was then designed to incorporate slots for bone cutting, holes for drilling of fixation points, and an optimised geometry which ensure ease of placement and use. The final device was then manufactured using additive manufacturing, to accurately replicate the complex surface topography and design features. To validate the design, the target patient anatomy was replicated using additive manufacturing and a ‘mock’ surgery was performed to assess the device performance.

To validate the design, the target patient anatomy was replicated using additive manufacturing and a ‘mock’ surgery was performed to assess the device performance. We found our design allowed for efficient placement and use during the mock surgery, confirming the potential of the devised process as a robust methodology for clinical implementation.1 IntroductionThe necessity to derive solutions for the enhancement and streamlining of surgical procedures are a vitally important goal towards achieving improved patient outcomes. From the surgeons perspective such solutions would comprise devices to assist inthe surgical resection process to minimise the removal of healthy bone from a patient, alongside the use of patient tailored, custom solutions to address an underlying condition (e.g. bone replacement implants).

bone replacement implants). In response to these demands, and coinciding with innovations in the fields of material science, advance design and additive manufacturing, devices are now emerging which address these criteria through the so called patient specific paradigm. Using such technologies, many innovative medical device innovations have been demonstrated with applications ranging from medical modelling [1-3], implantable devices [4, 5], prosthetics[6, 7], and orthotics [8, 9].With respect to surgical interventions, patient specific technologies are a rapidly emerging area in pre-clinical planning, bone resection and implantable devices with demonstrated applications in osteotomies [10], cancer resection [11, 12] and bone replacement [4, 5, 13], amongst others procedures.

Osteosarcoma surgeries (removal of bone tumours) can be a somewhat involved process requiring the use of chemotherapeutic action [14, 15], alongside the resection of the offending bone tissue [15, 16] in an attempt to completely remove the compromised tissues and halt any form of2398metastatic progression. Once performed, the residual uncompromised bone also requires fixation into its original orientation using either donor bone tissue or a load bearing implant alongside a fixation plate in order for the patient to regain normal use of the anatomy in question. Given the necessity to remove sections of corrupt bone from a patient, such procedures could significantly benefit from the use of patient specific technologies to streamline resection and piloting of drill holes for the fixation plate.

Additionally, given the unique topography of a patient’s anatomy, it would be desirable to move beyond the typical, generically sized fixation plates, to ones that more accurately mirror the patient’s bone topography, thereby realising a better fit amd improving the ability to retain the original orientation of the resected bone. Indeed there is significant potential for the fixation plate to be used in conjunction with a custom resection guide as part of a more comprehensive treatment option.In this study, we aim to develop a platform by which advanced design and 3D printing concepts can be used as a methodology to examine complex osteosarcoma in a simulated patient case study. Custom resection technology for such procedures have been demonstrated previously [11, 17] but have been relatively crude in their design and could significantly benefit from further innovations.

Custom resection technology for such procedures have been demonstrated previously [11, 17] but have been relatively crude in their design and could significantly benefit from further innovations. Initially, we replicate a patient case study who presented with a parosteal osteosarcoma of the radius, and replicate the tumour morphology on a publically available patient representative model, derived from CT scans. The advantage of this approach is that we can simulate the particular case study in a platform that is free from typical ethical constraints when using patient specific data yet still be reflective of a working example for examining the logistics of device development. We believe our demonstration to be the first example of such an approach.

We believe our demonstration to be the first example of such an approach. Using the devised patient representative model, we develop a custom resection guide which integrates functionality to avoid collision with the osteosarcoma, makes use of natural surface contours to provide optimal placement, integrates resection/drill/pin openings and a handle to easily manoeuvre the guide. We believe our guide to be the most sophisticated example of a patient specific guide found in the literature to address surgical intervention of distal radius osteosarcoma resection. The guide fundamentally realises all features required to perform all surgical intervention elements prior to fixation plate and implant/bone donor bone placement. To take the work a stage further, we also develop a patient specific fixation plate and mock, scaffold implant intended for use with the guide, providing a complete system for surgical intervention.

To take the work a stage further, we also develop a patient specific fixation plate and mock, scaffold implant intended for use with the guide, providing a complete system for surgical intervention. In the present study, we focus our attention on the design and product prototyping elements to demonstrate the potential feasibility of our approach. Ultimately, our methodology provides a relatively low cost means of examining the logistics of the device development under simulated circumstances, whereby the anatomical features of a patient are reproduced accurately, at low-cost and with no impact on the patient’s wellbeing.

We believe there to be significant potential for the devised devices to be 3D printed in biocompatible material for use in actual clinical circumstances.2 Methodology2.1 Anatomical data modellingModels of the patient data were constructed using anatomical data sets derived from CT scan data, based on the standardised Digital Imaging and Communications in Medicine (DICOM) file type with an image slice thickness of approximately 0.6mm. To convert the data into a working 3D digital model Mimics 19.0 (Materialise, Belgium) was used to isolate and extract the bone data based on a2399Hounsfield Unit (HU) threshold of 226-3071.

To convert the data into a working 3D digital model Mimics 19.0 (Materialise, Belgium) was used to isolate and extract the bone data based on a2399Hounsfield Unit (HU) threshold of 226-3071. Within Mimics bone models for the hand, wrist, ulna and radius were segmented from the wider data set and exported as STL files for further post processing.2.2 Cancer DesignThe segmented bone STL models were exported into 3-Matic Research 11.0 (Materialise, Belgium) for additional post processing to remove errors and major defects from the Mimics conversion process. To avoid ethical issues relating to the use of clinical patient data, we appealed to a published case study of a parosteal osteosarcoma of the radius [18], which we digitally reproduced as our example case for design of the resection guide (Figure 1). In this patient case, the tumour was defined as being attached to the surface, and wrapped around the bone.

In this patient case, the tumour was defined as being attached to the surface, and wrapped around the bone. CT scan images revealed the outer geometry of the tumour from several reference viewpoints. The written and visual evidence of the osteosarcoma in the patient study provided sufficient information to approximate the geometry of the tumour and to form our hypothetical digital tumour model for examination.2.3 Resection Guide DesignThe combined model of the patient specific anatomy of the arm and the tumour model were used as the template for the resection guide design, which was developed using 3-Matic. To construct the primary features of the resection guide, we manually designed elements in CAD software (Solidworks, DassaultSystemes, France). These features were then integrated onto a surface projection of the patient bone model, adjacent to the tumour site.

These features were then integrated onto a surface projection of the patient bone model, adjacent to the tumour site. Using this approach, the patient specific bone projection segments would allow for the ideal placement of the guide and self alignment using the natural contours of the bone as a locator guide. The use of traditional CAD, allows for the precise design of the bulk portions of the resections guide, such as the cutting slots, drill holes, Kirschner wires fixation ports and an ergonomic handle for manoeuvring/placement.When designing the resection guide consideration needs to be made as to the thickness of the part to ensure sufficient rigidity, but without enlarging the device to a point which results in complications during placement. Rigidity is important to ensure the device does not break during use and that flex it minimised to retain geometrical accuracy.

Rigidity is important to ensure the device does not break during use and that flex it minimised to retain geometrical accuracy. Keeping the device footprint to a minimum is important as to improve the patient recovery and avoid complications resulting from infection, general practise is to minimise the area exposed during a surgical intervention. Therefore the smaller the device the smaller the area that needs to potentially be exposed during surgery. As it is unknown a priori what the ideal geometrical attributes of the device should be, we examined various design iterations until a satisfactory candidate design was achieved, which was then manufactured for virtual surgery.2.4 Implant DesignIn this study we aimed to construct a concept of a patient specific bone fixation plate that could be used in conjunction with the resection guide. To achieve this we built a plate based on the typical design of a standard distal radius surgical plate [19].

To achieve this we built a plate based on the typical design of a standard distal radius surgical plate [19]. The plate is designed in 3-matic based on the surface topography of the mock patient's bone model and around the planned drill holes on the resection guide. Using this methodology we hoped to demonstrate2400a precise fit of the final bone fixation plate and alignment of the planned pilot holes. Additionally, as the system has been designed around the original positioning of the bone, once the plate is in position, we anticipate the remaining bone segments to return to their original position. For simplicity in the initial design, screw holes were designed to be non- locking, for use with fully threaded screws, and with an upper tapering section to allow the screws to sit flush with the mock plate.Alongside the fixation plate, we design a conceptual scaffold type implant to be used in place of donor bone material.

The scaffolds would potentially be representative of titanium implant that could allow for osseointegration or some form of bioprinted device. For now our intension of this element is purely cosmetic and for illustrative purposes, and is not intended as a developed treatment component. In future work we aim to develop this component further.Figure 01: a) Illustration of the surface osteosarcoma model formation process, b) a comparison of the CT scan image of the actual patient case study [18] and the digital representative model of the same case study and c) a concept view of the osteosarcoma on the patient.24012.5 Prototype manufacturingThis study focused on the design and surgical device developments facet. As they will not be used for direct clinical use, the various design iterations of the anatomical features, the resection guide and implant were realised using Filament Fused Fabrication (FFF) printing in commercially available ABS plastic.

This material provides a good representation of the final medical devices, alongside the patient’s representative anatomy. Evaluation prints were performed using a commercially available desktop printer (Flash Forge Dreamer, USA), which allow for a rapid turnaround time for part production. It is anticipated that the final resection guide will likely be printed using either FFF or Polyjet printing methodology in biocompatible materials (PC-ISO, MED 610, etc). However for evaluation purposes, the current ABS models would be comparable to parts made in such materials.3 Results3.1 Cancer Anatomical ModellingThe cancerous model was constructed by directly manipulating the patient specific radius bone model to shape the tumour.

Initially, a template portion of the distal region of the radius in the approximate location of the tumour, was segmented from the wider radius model of the patient and enlarged in the x-y-z direction by a scale factor of 2, yielding an enlargement of ΔX=25.6mm, ΔY=12.5mm and ΔZ=32.0mm. The resulting section was hollowed performing a Boolean subtraction of the original radius model, and trimmed to match the length of the template segmented region. The model was then segmented into two equal halves along the coronal plane, before being manually extruded to approximate the size of the tumour in all spatial dimensions. As we were not privy to the original patients medical imaging scans, we based the overall shape of the tumour model based on the relative sizes of the x-ray images of the presented case study [18] and by appeal to the typical gross specimen of a resected parosteal osteosarcoma [20].

Figure 1a), illustrates the template segment approach used to form the tumour, alongside the final model on the radius bone. Figure 1b) shows a direct comparison of the patient CT scan and the developed 3D model. Despite the differences in the patient anatomy, the general shape and location of the osteosarcoma has been accurately replicated and would be representative of an actual patient case. Figure 1c) finally shows a representative model of the bone, musculature and tumour as it is likely to be observed.3.2 Resection Guide DevelopmentResection of the compromised sections of bone can require the removal of between 6 to 20mm of ‘normal’ soft tissue, to ensure the complete removal of a tumour, thereby avoiding recurrence [21]. In the process of limb salvage surgical strategies it is important to minimise the removal of health bone tissue, to allow the best platform for functionality recovery.

In the process of limb salvage surgical strategies it is important to minimise the removal of health bone tissue, to allow the best platform for functionality recovery. However, a defined consensus has yet to be determined as to the minimal amount of healthy bone removal to ensure local non-recurrence of tumours. This lack of consensus has implications upon the design choices of the given resection guide design, and for this study, we optimised our design around a 10mm tolerance, which would be a good compromise between the ensuring tumour removal while minimising the removal of health bone.2402For the present case study, the resection guide will need to be designed to allow for easy fit to the patient's bone but with sufficient clearances to avoid issues with placement over the surface osteosarcoma, which protrudes outward from the bone.

We therefore examined an approach of using a surface projection of the bone surface, which will form the basis of the surface locator portion of the resection guide. Figure 2a) illustrates the design process by which the template bone surface data is constructed and manipulated to form the guide. The projection process comprises the initial isolation of a template region of bone, which is approximately 45-50mm from each extremity of the tumour. This region is then enlarged in all directions by 3mm, before a Boolean subtraction of the original bone structure, to create a hollow model which encompasses the whole topography of the template bone section. This is then digitally trimmed, firstly to open the hollow model at each extremity before divided into two halves at the upper and lower section of the model.

This is then digitally trimmed, firstly to open the hollow model at each extremity before divided into two halves at the upper and lower section of the model. The lower section is then discarded and a segment approximately 4-5mm either side of the tumour extremities is removed, resulting in the two halves of the resection guide surface locator model. The model is then trimmed at the distal region of the model to avoid placement issues at the interface with the wrist. Developing the guide to use the patient’s natural bone surface topography over the selected regions of the radius introduces natural markers for the positioning of the final resection guide, which is based on both the thickness of the bone at each half of the projected bone surface, and the vertical contours of the distal region of the bone. Therefore the guide would only have one ideal placement orientation whereby there was no movement in the guide once in position.

Therefore the guide would only have one ideal placement orientation whereby there was no movement in the guide once in position. This methodology ensures the geometrical accuracy of the pinning, drilling and resection features of the guide. It is also envisaged that to a minor degree, the protrusion of the tumour would provide further insight as to the placement of the guide based on the designed raised section of the guide.Figure 02: a) Illustration of patient specific bone surface template formation (purple) and b) annotated diagrams of the final resection guide concept.2403Following projection of the patient bone surfaces, a classical CAD approach was used to construct the drill/Kirschner wires holes and the resection slots. To aid with the design parameters, we appealed to standard instruments that would be used in an orthopaedic setting and which would be suitable given the size of the radius bone.

To aid with the design parameters, we appealed to standard instruments that would be used in an orthopaedic setting and which would be suitable given the size of the radius bone. The first constraint of the resection slot was to be placed 10mm away from each side on the tumour. The maximum width of the radius was digitally measured to be approximately 23mm, therefore opted to design slots suitable for a standard 20mm sagittal blade, with a blade thickness of 1.27mm (Stryker, USA). The slots comprised a rounded rectangular configuration, with a total internal length of 28mm and internal width of 1.8mm, as can be seen in figure 2b). This geometry was selected to ensure that there was sufficient clearances to encompass the full width of the radius bone, while also allowing enough space for the excursion width of the blade (22mm). The slots were also placed to be orthogonal to the bone at their location and made with a thickness of 4mm.

The slots were also placed to be orthogonal to the bone at their location and made with a thickness of 4mm. To complete the structural components, a bridge section over the osteosarcoma was developed, which raised the guide a height of 25mm above the height of the cutting slots. The bridge section also integrated two arcs, for finger handling of the guide by a user. The bridge section was initially made to a thickness of 3mm, however was found to have considerable flex, to a point where aggressive handling led to breakages. We therefore opted to increase the thickness to 6mm, which considerably increased the rigidity of the final model, to a point where flex was completely eliminated.

We therefore opted to increase the thickness to 6mm, which considerably increased the rigidity of the final model, to a point where flex was completely eliminated. In future development, we hope to characterise the tensile strength of the model and for now focus on the design for additive manufacturing elements of the guide.Figure 03: a) Design process for fixation plate development, b) design process for implant scaffold development and c) final digital patient model with fixation plate and implant in place.2404From direct discussion with orthopaedic surgeons, it was discovered that for a typical radius fixation plate ideally requires a minimum of 3 fixation points at either side of the resected bone to provide adequate support. In light of this, we incorporated this as part of our design constraints for both the resection guide and implant plate. For a typical surgical procedure on the distal radius, 2.4mm surgical locking screws can be used for fixation [10].

For a typical surgical procedure on the distal radius, 2.4mm surgical locking screws can be used for fixation [10]. We therefore designed the drill holes to have an internal diameter of 2.5mm and outer diameter of 5.5mm, which allowed for easy access of a standard surgical drill guide and the use of a 1.8mm drill bit to form a pilot hole for the screws. The requirement of designing the drill holes to be used with a standard metallic surgical drill guide as opposed to simply using the 3D printed guide eliminates any complications that may occur due to the drill damaging the guide and releasing debris in the surrounding tissue. The drill holes were located long the centre line of the radius, relative to the resection slots and at a regular spacing of approximately 5mm from the outer diameters of either the resection slots or drill holes. This spacing allowed for adequate clearances so that the screws would not intersect each other when placed in the bone.

This spacing allowed for adequate clearances so that the screws would not intersect each other when placed in the bone. To confirm this, a ‘digital screw’ was created and used to guide the placement of the drill holes. Finally, Kirschner wires holes were created at either extremity of the guide to ensure the guide could be temporarily pinned in position during resection. Kirschner wires can be used as an effective methodology to hold resection templates in place during surgery [10, 12] and are an important feature to retain the geometrical accuracy of the guide during use. Wire holes with an inner diameter of 1.5mm and outer diameter of 4.5mm were designed based on standard Kirschner wire sizes, and were angled such that when the pins are in position the tension acts to lock the guide in place.

The overall final guide concept can be seen in figure 1b).3.3 Fixation Plate and Implant ConceptTo complement the patient specific guide, a concept implant was developed to both replace the resected bone and to provide support to the segmented bone sections. As with the resection guide, these components were designed to be patient specific and match the original anatomical features of the patient. Figure 3a) illustrates the design process for construction of the fixation plate. Initially, using the bone projection that was used to form the base of the guide, the basic form of the plate is realised. As the resection guide had already incorporated the drill holes in the design, we could use the centre point of the holes as a template for the screw holes on the plate. As the final design would likely be manufactured using Ti-64, it is important to reduce the overall mass of the part to a minimum.

As the final design would likely be manufactured using Ti-64, it is important to reduce the overall mass of the part to a minimum. Therefore, the template systematically underwent topology optimisation to reduce the overall mass of the plate to realise the final design as seen in Figure 3a).The implant to replace the resected bone will comprise an open scaffolds structure. We have previously reported the rapid development of scaffold structureswhich can be applied in titanium implantable devices for light weighting [4]. However, in this study the scaffold design is primarily for conceptualisation purposes, but potentially could be applied to next generation bioprinted scaffolds or titanium implant as the technologies mature. Briefly, the process for scaffold formation involves the segmentation of the bone template, before applying a fixed geometry unit cell structure throughout the digital model, as described previously [22].

In this instance we opted for a Gyroid lattice unit cell structure, with a unit cell size of 3x3x3mm, and the final scaffold structure can be seen in figure 3b). Finally, a virtual representation of the implant, fixation plate and patient anatomy was constructed to better visualise the parts in situ, and can be seen in figure 3c).24053.4 Mock SurgeryTo qualitatively test the efficacy of the devised resection guide, the implant and fixation plate, a mock surgery was performed, whereby the resection guide was pinned onto a 3D printed model of the patient’s arm and osteosarcoma as can be seen in figure 4ai). As we did not have access to a sagittal saw, we opted to mark the region of the cut using a thin blade, and the model was cut using a circulating saw. Several repeat cuts were performed and the variance in the cutting position from the midpoint of the cutting guide varied by approximately ±0.6mm.

Several repeat cuts were performed and the variance in the cutting position from the midpoint of the cutting guide varied by approximately ±0.6mm. In addition to the cuts, a standard 1.8mm drill bit was used to drill pilot holes in the model to attachment of the screws into the fixation plate concept. It was found that the drill holes were achieved to an accuracy of ±0.3mm of the intended centre point. Once all mock resection and drill piloting had been performed the pins and guide were removed and the implant and plate were screwed into position. The entire procedure from start to finish took approximately 2-3 minutes to perform, reflecting the potential time it would take the surgeon for this portion of the procedure within theatre.Figure 4a) shows the stages of the model, pre and post mock resection. It was found that the implant fix accurately into the bone model, along with the placement of the screwed fixation plate.

It was found that the implant fix accurately into the bone model, along with the placement of the screwed fixation plate. Equally, the resection guide had sufficient rigidity to allow for aggressive handling when simulating the placement of Kirschner wires, marking cuts and drilling of the pilot holes. It was noted that there was some ability for the screws to deviate from their desired directional placement, which we believe is due to the constraints of the model being fabricated in ABS plastic as opposed to real bone. It is believed that if the procedure is performed on actual bone that the rigidity of the bone would allow for greater directional conformity of fixation screws.

It is believed that if the procedure is performed on actual bone that the rigidity of the bone would allow for greater directional conformity of fixation screws. Overall, we are confident that there is considerable potential of our approach for development as a clinical treatment option, should the resection guide be manufactured in biocompatible materials and the implant/fixation plate in Titanium 64, the current benchmark for titanium implants.Figure 04: a) 3D printed hand, wrist, ulna and radius models with i) surface osteosarcoma (black), ii) the osteosarcoma and resection guide concept in situ and iii) the resected radius with fixation plate and scaffold implant.

b) close up views of the radius with i) the resection guide pinned in place and ii) the fixation plate and scaffold concept.24064 ConclusionIn this study we present our initial findings in the development of a resection guide and implantable treatment concepts, which could serve as methodologies for the treatment of a surface osteosarcoma on the distal radius. The guide was found to accurately match the contours of the mock patients bone anatomy and provided a high degree of efficacy to minimise the removal of potential health bone tissues, with resection tolerances being approximately ±0.6mm from the desired location. We also demonstrate a patient specific fixation plate and mock implant, which conform very well to the bone topography, providing efficacious fixation and reproduction of the patient’s original bone anatomy.

It is our hope that the methodology outlined in this preliminary study would provide guidance to professionals seeking to develop a robust design methodology for resection guide and fixation plate development and which could be applied to a range of different clinical case studies.AcknowledgementsThis research was conducted by the Australian Research Council Industrial Transformation Training Centre in Additive Biomanufacturing (IC160100026). All resources for the study were provided courtesy of the School of Engineering at Deakin University.References1.2.3.4.5.6.7.8.9.10.11.12.Salmi, M., et al., Accuracy of medical models made by additive manufacturing (rapid manufacturing). Journal of Cranio-Maxillofacial Surgery, 2013. 41(7): p. 603-609. Mitsouras, D., et al., Medical 3D Printing for the Radiologist. RadioGraphics, 2015. 35(7): p. 1965-1988.

velopment of Direct SLS Processing for Production of Cermet Composite Turbine Sealing Components - Part I T. Fuestingt, L. Brownt, S. Das:t, N. Harlan:t, G. Lee:t, J. J. Beaman*, D. L. Bourell*, J. W. Barlow*, K. Sargent'Abstractto date ofSLS (selective laser sintering) technologies for production of cermet composite turbine sealing components, the particular application being an abrasive blade tip. The component chosen for the application is an integral part of the low pressure turbine in a IHP'TET (Integrated High Performance Turbine Engine Technology) demonstrator engine. Both indirect and direct SLS techniques are being developed. Initial trials and process development involved the use of fugitive polymeric binders. Sequential refinements were performed to develop a binderless direct SLS process.

Initial trials and process development involved the use of fugitive polymeric binders. Sequential refinements were performed to develop a binderless direct SLS process. Results from mechanical testing indicate that acceptable microstructure and properties are attainable by SLS with substantial cost savings as compared to the currently employed production method. This is the rust instance of direct SFF methods applied to a functional componentThis paper presents the developmentINTRODUCTIONSince its inception in the 1930's, the gas turbine engine has grown to be the workhorse power plant of modem aircraft. Advanced technology propulsion systems are very important from an overall system affordability viewpoint since higher engine performance can result in a smaller airframe.

Advanced technology propulsion systems are very important from an overall system affordability viewpoint since higher engine performance can result in a smaller airframe. In 1988, the United States Department of Defense and the National Aeronautics and Space Administration combined forces with industry to focus gas turbine engine research and development efforts toward the goal of doubling current levels of propulsion capability by the IHPTET (Integrated High Performance Turbine Engine turn-of-the-century. Known as the goals are to Technology) program, increase thrust-to-weight, reduce fuel burn, and reduce the cost of high performance gas turbine engines.this aggressive technology development plan's specificIncreased affordability can be accomplished by either increasing engine performance without cost Penalty or by decreasing engine cost without sacrificing performance.

The two principal aspects that IHPTET is targeting to reduce engine life cycle cost are engine unit acquisition cost and, operating and support costs. The technological developments required to lower the cost of future engines are advanced materials, innovative structural designs, improved aerothermodYfiamics, advanced computational methods, and advanced manufacturing techniques. The integration of these advanced technologies will not only provide the benefit of improved performance, but will also help to make engine systems more robust and more affordable. IHPTET cost reduction studies! have found that a key to engine cost reduction is cost effective fabrication and manufacturing.

IHPTET cost reduction studies! have found that a key to engine cost reduction is cost effective fabrication and manufacturing. These studies have also shown that several principal factors can contribute to lower manufacturing costs for the advanced technology engine, including its smaller size, reduced parts count and cost effective manufacturing methods.An example of advanced manufacturing methods that can impact cost reduction of engine components is SFF. This technology is striving towards decoupling cost from volume2 by making it possible to produce the first unit at a recurring cost equal to the hundredth unie. The IHPTET It is a cermet composite program has chosen an application to demonstrate this technology.

The IHPTET It is a cermet composite program has chosen an application to demonstrate this technology. component that is an integral part of the low pressure turbine in an IHPTET demonstrator engine.t Allison Engine Company, Indianapolis, Indiana t Laboratory for Freeform Fabricationy University of Texas at Austin § Wright Laboratory, Wright-Patterson AFB, Ohio39This piper begins with a brief background on gas turbine sealing technology. The objectives for pursuing the research described herein are discussed next. The potential for significant cost savings by pursuing SLS as a manufacturing technique is demonstrated. The Allison EngiDe Co. "standard cermet" production technique is described, followed by a description of materials and methods used in this research. The results section compares properties ofoennet material processed by SLS with those of cermet produced by the standard production tecbDique.

The results section compares properties ofoennet material processed by SLS with those of cermet produced by the standard production tecbDique. Problems encountered with materials interaction in the cermet composite and their sol1l1ioDs are discussed. Conclusions and plans for future work are presented.BACKGROUNDGas turbine efficiency is highly dependent upon minimizing leakage of the gas away from and around the working gas path. Therefore, clearance between rotating and static parts is critical. This (clearance changes with component expansion and contraction due to the thermal cycling experienced in gas turbines. One of the primary methods developed for accounting for this expansion and contraction is an abradable seal system4 .5. This seal system works by attaching a cermet composite abrasive blade tip (ABT) to the sealing surface of the rotating component.

This seal system works by attaching a cermet composite abrasive blade tip (ABT) to the sealing surface of the rotating component. The circular stationary component is coated with an appropriately engineered semi-porous abradable ceramic which is abraded by the rotating cermet to form a shroud tyPe gas path seal around the rotating component.The semi-porous ceramic component is relatively inexpensive and simple to manufacture. However, the abrasive cermet (0.035-0.060 inches thickness) is difficult to manufacture on a production scale and contains expensive components. The cost of the raw materials makes scrap highly undesirable. Manufacturing of this layer typically involves rolling and pressing a mixture of the ceramic and metal alloy powders with an organic binder to the correct thickness. The rolled mat is then vacuum sintered. It is evaluated for proper microstructure, machine ground to proper thickness, laser cut and metallurgically bonded to the rotating hardware.

The rolled mat is then vacuum sintered. It is evaluated for proper microstructure, machine ground to proper thickness, laser cut and metallurgically bonded to the rotating hardware. This complete process is labor intensive and can result in a significant amount of scrap if stringent manufacturing processes are not adhered to.Due to the difficulties encountered with standard abrasive cermet production, an alternate approach for producing this material is desired. Since the standard production material is capable of being laser cut to final shape, laser sintering of the constituent powders was the primary alternate approach chosen. SLS (Selective Laser Sintering) is a rapid manufacturing technique that builds parts from a powder base. SLS creates freeform 3-dimensional solid objects without the use of any tooling or human intervention.

SLS creates freeform 3-dimensional solid objects without the use of any tooling or human intervention. Details on this process are described elsewhere6SLS has been commercially developed by DTM Corporation and is in use for prototyping models from polymeric powders and for creating investment casting wax patterns. More recently, DTM has introduced RapidTool, a technology for creating prototype metal tooling from polymer • This process uses indirect SLS techniques lO which involve the use of coated metal powders9 fugitive polymer binders to produce a "green" shape that is subsequently post-processed by binder bum-out, sintering and infiltration to produce high density metal and ceramic parts and tooling11,12,13,14,15. A functional metal part, porous or fully dense, has not been made previously by direct SLS.

A functional metal part, porous or fully dense, has not been made previously by direct SLS. The terminology "direct SLS" implies a binderless SLS process in which the material constituents are directly laser sintered to produce a high density part requiring little or no post-processing.The expensive and difficult to manufacture abrasive cermet composite can potentially be40,7,8.directly laser sintered, therefore making it an ideal part for processing by direct SLS. An SLS process for producing abrasive cermet composite can potentially provide a uniform, repeatable manufacturing process that would eliminate much of the manual labor as well as reduce scrap and make all unsintered fall off materials reusable.OBJECTIVESThe primary goal of research described herein is to develop an SLS process capable of producing the cermet. The cermet produced by SLS must have properties equivalent to the current production material used by Allison Engine Company.

The cermet produced by SLS must have properties equivalent to the current production material used by Allison Engine Company. The process developed should allow for possible reuse of all unsintered fall off material. Another goal is to eliminate a low melting point braze constituent which will extend oxidation life and possibly allow higher operating temperatures.Reducing production costs and achieving a repeatable production process are the primary motivation for pursuing SLS as a manufacturing technique in this application. The typical manufacturing method of tape casting and rolling requires a fixed amount of material input per lot, independent of the actual parts needed per lot size, up to lot sizes producing 1000 parts. For prototype lot sizes to produce 100 parts or less, this method becomes expensive considering the cost of the materials involved. Moreover, sintered material left over after the laser cutting step, is scrap.

Moreover, sintered material left over after the laser cutting step, is scrap. Since direct SLS does not utilize a binder, it eliminates nearly all scrap "fall-off', material by enabling recycling of all fall- off materials from a lot. This significantly reduces the cost and results in the primary portion of the cost savings. The remainder of the cost savings are realized by elimination of labor intensive tape casting and rolling in the currently employed production process. A cost break-down, shown in Figure 1, was performed to illustrate the potential benefits of developing an SLS process. The details and assumptions used to formulate this break down are provided in Table 1. Obtaining a repeatable production process also results in very large cost savings which are not directly measurable. The current production process has variable yield, resulting in a significant amount of scrap due to unacceptable porosity.ABT Manufacturing Cost Analysis10ססoo1ססoo~ a.

Cost breakdownMATERIALS AND METHODS Standard Cermet ProductionThe standard abrasive cermet composite is composed of two different types of titanium coated ceramic abrasive grit, a nickel alloy matrix, and a lower melting point cobalt based braze material. The compositions of the metal alloy components are provided in Table 2. The titanium coating on the ceramic grit functions as a bonding interlayer to promote a metallurgical bond as opposed to a mechanical interlocking bond which would result from using uncoated grit. A metallurgical bond is more effective in preventing grit pull out which occurs during rubbing or abrading of the cermet composite. Physically or mechanically interlocked grit has a high tendency to be pulled out of the matrix by the shear forces encountered during rubbing.

Physically or mechanically interlocked grit has a high tendency to be pulled out of the matrix by the shear forces encountered during rubbing. Removal of grit (1) the particles results in reduced abrasive for cutting a clear path which has two consequences: metallic matrix will be rubbing the ceramic abradable seal material, increasing friction forces and reducing efficiency, and (2) occurring but less likely, the pullout introduces a gap for compressed gas to leak around the working path. Additional details are described elsewhere16,17.Material Mar-M-247 *CMSX-3 0 Amdry788Ni Bal Bal 21.0 Bal * CMSX-3 is a re.e:istered trademark of Cannon-Muskegon Corp. o Amdry 788 is braze alloy from Sulzer Plasma Technik, Inc. Table 2.

o Amdry 788 is braze alloy from Sulzer Plasma Technik, Inc. Table 2. Nominal compositions of matrix materials used in the cermetsCo 10.0 4.6Cr 8.4 7.8 22.0AI 5.5 5.6Ti 1.0 1.0W 10.0 8.0 14.0SiB 0.012.02.0Mo 0.7 0.5Ta 3.1 6.0To produce a "lot" of this material via tape casting and rolling, a fixed amount of material input is required regardless of the lot size. These materials are mixed by low energy ball milling followed by the addition of an organic binder. The mixture is then tape cast and subsequently rolled to eliminate porosity and obtain the desired thickness. After rolling, the material is fired in a vacuum furnace to bum off the binder and consolidate the mixture by melting and diffusing the braze constituent. After densification the material is metallographically analyzed to determine whether it meets homogeneity and porosity requirements.

After densification the material is metallographically analyzed to determine whether it meets homogeneity and porosity requirements. To meet homogeneity requirements,42Hf 1.4 0.1the sintered material must exhibit no discrete layering or agglomeration of abrasive particles which would result in grit pullout causing a decrease in efficiency. Porosity requirements allow for no more than 20% linear porosity as determined from a metallographic montage using a line intercept method. If the material meets these specifications, coupons are removed and brazed into the gage section of a test bar conforming to ASTM E139 and subjected to an application specific stress rupture test. The details of this test are provided in the results section of this paper.The strength and oxidation resistance of the cobalt braze component are factors currently limiting the application temperatures.

There are alternate methods of making brazeless cermets, such as HIP (hot isostatic pressing), ROC (rapid omni-directional compaction) and Electro depostion. However, they are more labor intensive. Therefore, development of a direct SLS process for production of brazeless cermets will not only provide significant cost savings but also performance enhancement over that of the braze containing composition.Cermet Production by SLS ApparatusThe apparatus used for this research included SLS workstations equipped with Nd:YAG and CO2 lasers. The indirect SLS work was done at ambient atmosphere using a 25 Watt CO2 laser. The direct SLS work was conducted on SLS workstations equipped with 100 Watt and 250 Watt Nd:YAG lasers, powder bed heating capability and controlled atmosphere.Processing MethodsTwo SLS processing methods were investigated concurrently.

The most desired method is direct SLS which involves directly melting and consolidating selected regions of a powder bed to form a desired shape having high or full density. Direct metal laser sintering involves melting the component matrix and obtaining the appropriate amount of flow from the molten material. The appropriate amount of flow is critical and can be described as the flow that eliminates porosity, produces a highly dense part maintaining dimensional tolerances and minimizes other defects such as hot tearing. The appropriate amount of flow is controlled by several factors such as atmosphere, powder bed heat and three characteristic variables affecting laser energy density: laser power, scan spacing and scan speed.These SLS processing parameters can be refmed within a working window depending upon the mechanical properties desired. In addition, a post-process heat treatment may be applied to modify the microstructure.

In addition, a post-process heat treatment may be applied to modify the microstructure. The parameter used for microstructure refinement was the energy density18 defmed asp AN = v.~ (J/cm2)whereP is the incident laser power (Watts) v is the laser scan speed (cm/s) ~ is the scan spacing (cm)An alternative method investigated is indirect SLS. The term "indirect" implies that the metal powder is not melted directly by the laser. Instead, a low glass transition temperature polymer binder either mixed into the powder or present as a coating on the powder particles softens and flows during selective laser sintering. This polymer acts as the binding phase, giving43strength to the "green" part. The "green" part is post processed to bum off the polymer and sinter the metal powder.

This polymer acts as the binding phase, giving43strength to the "green" part. The "green" part is post processed to bum off the polymer and sinter the metal powder. The goal of this processing method was to develop the procedure into a process that incorporates laser volatilization of the binder immediately before or during direct SLS.Indirect SLS material preparationThe constituent components were combined according to the compositions shown in Table 3 below and ball milled in a tungsten carbide container. The composition of the cobalt braze was varied to determine its effect on the final properties of the sintered composite, providing an optimized composition with respect to braze content. The amount of cermet grit remained constant in all mixtures.

The amount of cermet grit remained constant in all mixtures. All grit except the grit used in Mixture 5 was coated with titanium.Material Composition Composition Composition Composition Composition CompositionAlloy Braze Grit Binder1 Bal 24.3 wt.% Ti Coated 2.0wt.%2 Bal 21 wt.% Ti Coated 2.0 wt.%3 Bal 18 wt.% Ti Coated 2.0wt.%4 Bal 15 wt.% Ti Coated 2.0 wt.%5 Bal 24.3 wt.% Uncoated 2.0 wt.%6 Bal 24.3 wt.% Ti Coated 2.0 wt.%Table 3.Indirect SLS compositionsInitial trials indicated that there was a sporadic problem with grit segregation. To eliminate this problem, a slurry was created by adding acetone. The slurry was mixed in a motor driven stirrer. During mixing, the acetone evaporated leaving a highly viscous slurry paste which was pressed to form a thin cake and dried to volatilize the remaining solvent. The cake was crushed using a mortar and pestle to form a coarse powder composed of roughly 500 Jlm particle agglomerates.

The cake was crushed using a mortar and pestle to form a coarse powder composed of roughly 500 Jlm particle agglomerates. SEM micrographs are provided in Figure 2.The agglomerate powder was placed in the stainless steel mold of a compression molding The platens of the molding machine were preheated above the binder softening press. temperature. Pressure was maintained during heat up and cool down. The fractional density of the bound and pressed preform was measured, using the linear intercept method, at 60%.Indirect SLS processingSLS processing included cutting the desired shape from the pressed preform using a 25 Watt CO2 laser. Several passes were necessary at this laser power to completely penetrate the 0.060" preform thickness. Use of higher laser power of the order of 100 Watts would facilitate cutting the shape in a single pass.

Use of higher laser power of the order of 100 Watts would facilitate cutting the shape in a single pass. Other process trials addressed the volatilization of the binder utilizing low power scanning to facilitate direct SLS processing.Direct SLS material preparationMaterial preparation for the direct SLS method was simpler. The metal alloy and ceramic constituents were combined according to mixture 3 shown in Table 3 above and blended for 4-6 hours prior to loading in the SLS chamber.44Figure 2. SEM micrographs of AgglomeratesDirect SLS processingTo characterize the behavior of the metallic constituents of this material system with respect to laser irradiation, the superalloy and the cobalt braze were individually processed by SLS. These initial screening trials indicated that structurally sound monolithic single layers could be produced. Typical microstructures from these trials are shown in Figures 3 through 5.

These initial screening trials indicated that structurally sound monolithic single layers could be produced. Typical microstructures from these trials are shown in Figures 3 through 5. However when the metal alloys were blended together the material exhibited severe balling. The balling of the alloy mixture is demonstrated by the photomicrograph of the sample in Figure 5. The uniform dendritic microstructure is indicative of thorough melting and mixing.Figure 3. Direct SLS Mar-M-247, no etch, magnification 50X45Figure 4. Direct SLS cobalt braze, no etch, magnification 200 XFigure S. Direct SLS Mar-M-247 & braze alloy, magnification 200 XAs stated earlier, one of the goals of this research was to investigate the elimination of the cobalt braze alloy from the material system. Due to the severe balling exhibited by the standard mixture recipe, direct SLS of a brazeless cermet composition was pursued.

Due to the severe balling exhibited by the standard mixture recipe, direct SLS of a brazeless cermet composition was pursued. This composition was formulated by replacing the weight fraction of the cobalt braze by the nickel base superalloy. This composition was utilized in all direct SLS experiments conducted for the remainder of the program.46.

The association’s mission is twofold; to develop metal Additive Manufacturing technology and to develop binder jetting-type machine for the rapid production of sand molds. The goal of TRAFAM project is the development of innovative metal AM systems that will meet the world’s highest standards and the development of AM technology for high value-added products for aerospace, medical, and transportation industries etc. In this paper, the latest development status of the project, especially the development status of the electron beam PBF type machines, the laser PBF type machine and the laser DED type machines, is introduced. In addition, the numerical simulation software for metal AM based on multi-scale and multi- physics model is introduced.IntroductionThe Ministry of Economy, Trade and Industry (METI) established a study group on New “Monodzukuri (manufacturing)” in October, 2013 because Japan lags behind the U.S.A. and Europe in addressing Additive Manufacturing (AM) technology.

and Europe in addressing Additive Manufacturing (AM) technology. The study group proposed the future directions of “Monodzukuri”, and finally compiled the results of the meetings into a report [1]. Also the study group offered the issues to tackle; (1) developing technology in which devices, materials and software, (2) developing environments, (3) fostering human resources, and (4) seeking ideal approaches to creating enterprises. On the basis of this offer, METI invested about $36.5 million to establish a new research association, in FY2014 Technology Research Association for Future Additive Manufacturing (TRAFAM), in order to develop metal AM systems and a binder jetting machine for producing sand molds.In this paper, the latest development status of the project, especially, the development status of the electron beam powder bed fusion (PBF) type machines, the laser PBF type machine and the laser directed energy deposition (DED) type machines is introduced.

In addition, the simulation software for metal AM based on multi-scale and multi-physics model developed by TRAFAM is introduced.The role of TRAFAMMETI invested about $36.5 million in FY2014 to launch a new research association, Technology Research Association for Future Additive Manufacturing (TRAFAM) [2], in order to develop metal AM systems and a binder jetting machine for producing sand molds rapidly. The members of TRAFAM include three academic institutions and 34 companies in FY2018 as shown in Fig.1. The mission of TRAFAM project is to establish a new manufacturing industry in Japan centering on AM systems that will give rise to the next generation of innovative products.

The goal of TRAFAM project [3] is the development of the innovative AM machines meeting the world’s highest standards as shown in Table 1;Enhanced speed (Approximately 10 times the speed at FY2013) - Enhanced accuracy (Approximately 5 times the accuracy at FY2013) - Upsizing (Approximately 3 times the molding range at FY2013) - Multi-material structure (Different types of metal materials can be used) 1641TRAFAM developed the PBF type and DED type of AM machines with electron beam or laser beam systems The software for controlling AM machines and the production technology of metal powder for AM technology are developed in order to make AM machines higher performance.Technology Research Association for Future Additive Manufacturing (TRAFAM)President: Atsushi Maekawa (Director, President, Mitsubishi Heavy Industries Forklift, Engine & Turbocharger Holdings, Lid.)1. Next generation industrial 3D printer technology project PL: Prof.

The second EB-PBF machine (Fig.2(b)) which is mainly used for TiAl alloy is equipped with an electron gun of LaB6 cathode and an electron optic system at 10 kW and a prevention system of electric discharge. Examples of trial products fabricated by the first machine are shown in Fig.3. JEOL is also developing a dispersing process technology of multi-material powders as shown Fig.4.1642(a) First machine(B) Second machineFig.2 EB-PBF machines developed by JEOL(a) Closed impeller (Early stage)(b) Opened impeller (middle stage)Fig.3 Examples of trial products by JEOL machinesCopper and M2 powder Accuracy of positioning: 200 μm Fig.4 Results of multi-material powder dispersion(2) Large–scale EB-PBF machineTADA ELECTRIC CO. LTD., which is one of MITSUBISHI ELECTRIC CO. group, developed a large- scale EB-PBF prototype machine as shown in Fig.5. The build size is 500x500x600 mm and the rated output is 6 kW by the original long lifetime cathode (over1000 h).

group, developed a large- scale EB-PBF prototype machine as shown in Fig.5. The build size is 500x500x600 mm and the rated output is 6 kW by the original long lifetime cathode (over1000 h). The high-speed deflection of electron beam control system was developed to control an electron beam on a large area of powder bed stably. Also, the powder recycle system for large-scale machine was developed as shown in Fig.6.1643Advanced thermal insulation systemHigh speed scanning/ High precision electron beam gunI,Continuous powderHigh speed squeeze/ Powder feed control' supply system (Back side)Removable mold boxFig.5 Large-scale EB-PBF prototype machine developed by TADA ELECTRICPowder pulverizing unitnPowder collection/convey unitFig.6 Powder recycle system for large-scale machine(3) Large-scale LB-PBF machineMATSUURA MACHINARY CO. developed a large-scale LB-PBF prototype machine (Fig.7(a)) and a large-scale LB-PBF second machine (Fig.7(b)).

developed a LB-DED prototype machine (Fig.9) equipped with a 6 kW laser and a high performance nozzle. This machine type is multi-layer and the build speed is 510 cc/h at maximum. An example of a complex-shaped pipe with accurate and smooth surface owing to not only a high performance nozzle but also the CAM software for 5-axis control & multi-material DED is shown in Fig.10.Fig.9 LB-DED multi-layer prototype machineFig.10 An example of complex-shapedpipe (INCONEL 718)A large-scale LB-DED prototype machine was developed as shown in Fig.11.

The specifications are as follows:Maximum dimension of built parts: 2000 mm in diameter x1000 mm in height - Laser power for building: 6 kWx2 - Number of axis: Head 3 axis (Y, Z, A), Table 2 axis (X, C) An example of a large fan case after machining is shown in Fig.12.Fig.11 Large-scale LB-DED prototype machine Fig.12 An example of fan case after machining(Size:(cid:557)710 x h400 x t10 mm, Black area is as-deposited.)1645(5) LB-DED and Milling hybrid-type machineMITUBISHI HEAVY INDUSTRIES MACHINE TOOL CO. LTD. developed a LB-DED and milling prototype machine (Fig.13) equipped with a 6 kW laser and a high performance nozzle. Examples of a hollow body fabricated on a shaft are shown in Fig.14. The build speed is 500 cc/h at 6 kW at maximum, and the powder is able to switch to different one. This machine is suitable for manufacturing larger parts, repairing and printing different powders or functionally graded materials.

This machine is suitable for manufacturing larger parts, repairing and printing different powders or functionally graded materials. The monitoring feedback of this system was developed so as to manufacture parts with high quality stably. The commercial machine is shown in Fig.15. The CAM software for 5-axis control & multi-material DED was developed in collaboration with the DED group as shown in Fig.16.Fig.13 LB-DED & Milling prototype machine Fig.14 Examples of hollow body fabricated on a shaft(SUS316L & INCONEL 718)Fig.15 LB-DED commercial machine Fig.16 CAM software for 5-axis control & multi-material DEDDevelopment of simulation software for metal AMThe development of numerical simulation is one of the most significant issues not only to manufacture high-quality products but also to realize “Digital Twin” for AM processe.

The numerical simulation software analyzing the melting and solidification phenomena, the prediction of microstructure and the thermal deformation has been developed on the basis of multi-scale modeling technique and multi-physics simulation. Figure 17 demonstrates the schematic illustration of multi-scale modeling strategy for the AM process in TRAFAM project.

The micro-simulation for melting-solidification analysis based on a two-fluid model and the thermal deformation simulation using inherent strain method were developed.1646Development of simulation software for metal AM based on multi-scale and multi-physics modelMacro-simulation for melting-solidification analysis using commercial softwareQJ s E l=msMicro-simulation for melting-solidification analysis based on two-fluid modelThermal deformation simulation using inherent strain methodµsPrediction of microstructure & making solidification map using commercial softwareµmmm Sizemustration of multi-scale modeling strategy for the metal AM process Fig.17 Schematic illustration of multi-scale modeling strategy for the metal AM process in TRAFAM projectThe numerical simulation software of melting and solidification phenomena in metal PBF process was developed on the basis of a two-fluid model.

The significant physical phenomena during melting and solidification such as keyhole, plume, and spatter generation due to interfacial pressure difference between gas and liquid phases can be treated using this software [4]. The gas-liquid two-phase flow is treated with the Eulerian-Eulerian two-fluid model. Figure 18 shows the top view of single track of Inconel 718 by simulation. The scattering of the spatter of melting metal as well as the melt pool are able to be observed. In the case of the one-fluid model [5], although the rolling waves of the molten metal surface can be expressed, the scattering of the spatter cannot be simulated. As a result, the simulated result is similar to the experimental one shown in Fig.18(b).Laser power (W)19781180 .,.0I 5 .,, ., 0 0 ~"'OilC -~ .; () Vl0 0""' N·-- .(a) Simulation ( ) Sil ti. -·~. I -·~ · ..... -,,- -·.,, ......... -.. .(b)(cid:3)Experimental result Fig.18 Melting and solidification phenomena.

-·~. I -·~ · ..... -,,- -·.,, ......... -.. .(b)(cid:3)Experimental result Fig.18 Melting and solidification phenomena. (a) Top view of single bead molding processes of Inconel 718 by simulation [4], (b) Experimental result of Inconel 718 observed using a high-speed cameraMacro-simulation for melting-solidification analysis using commercial software was performed and predicted fabrication conditions such as laser power, (cid:86)(cid:70)(cid:68)(cid:81)(cid:3)speed, hatch pitch and so on. The analysis of melting- 1647solidification phenomenon of a melt pool by a computational fluid dynamics (CFD) was performed, and the solidification maps, displaying solidification microstructure as a function of solidification rate and thermal gradient at the solidification front, were determined by using a cellular automaton simulation.In addition, the thermal deformation simulation software for LB-PBF, EB-PBF and DED using inherent strain method was developed.

The results analyzed by the inherent strain method coincided considerably with those of the thermal elastic-plastic FEM analysis by optimization of the inherent strain of material.SummaryTechnology Research Association for Future Additive Manufacturing (TRAFAM) was carried out the following two projects in order to develop the innovative Additive Manufacturing systems that would meet the world's highest standards and the manufacturing technologies for high-value-added products. The schematic chart of the Next-generation industrial 3D printers project is shown in Fig.19. The main results obtained are as follows:(A) Next-generation industrial 3D printers project(1) TRAFAM developed the PBF type and DED type of AM prototype machines with electron beam or laser beam systems. The software for controlling AM machines and the production technology of metal powder for AM technology were developed in order to make AM machines higher performance.

The software for controlling AM machines and the production technology of metal powder for AM technology were developed in order to make AM machines higher performance. (2) Morphology and temperature of a melt pool were observed and analyzed by using the electron-beam powder-bed fusion (EB-PBF) system, equipped with a high-speed camera. Based on this basic survey, the advanced monitoring system for the EB-PBF has also been designed and prototyped. (3) The electrical and thermal properties of alloy powders used in the EB-PBF system were measured, and the effect of a surface oxide layer was analyzed in order to consider the methodology to suppress the smoke phenomena in EB-PBF system.

(4) The analysis of melting and solidification phenomenon of a melt pool by a computational fluid dynamics (CFD), solidification maps, displaying solidification microstructure as a function of solidification rate and thermal gradient at the solidification front, were determined by using a cellular automaton simulation, not experiments. (5) Considering the melting and solidification phenomena by laser radiation, the macroscopic simulation was carried out by using the newly developed heat input model considering the different laser absorption factor values by the powder layer, solidified part, and liquid phase in the unsteady heat conduction analysis. And, by introducing the overlap ratio, it was suggested that the optimum value of energy density could be predicted.(6) The multi-physics model in the LB-PBF system was constructed to simulate the melting and solidification phenomena at a microscopic level by using a supercomputer.

(7) The thermal elastic-plastic simulation program was developed by using the inherent strain method in order to predict the deformation of the parts fabricated by DED, LB-PBF and EB-PBF. The results analyzed by the inherent strain method coincided considerably with those of the thermal elastic-plastic FEM analysis by optimization of the inherent strain of materials. (8) The tensile tests and fatigue tests were carried out using the specimens fabricated by PBF and DED type prototype machines. The tensile strength of the as-built and HIPed specimens was similar to that of the wrought materials. The fatigue strength of the as-built specimens was lower than that of the wrought materials, while the fatigue strength of the HIPed specimens was similar to that of the wrought materials.

The fatigue strength of the as-built specimens was lower than that of the wrought materials, while the fatigue strength of the HIPed specimens was similar to that of the wrought materials. (9) The obtained data have been stored in the database designed and developed in this project.1648UsersSchematic chart of TRAFAM projectFig.19 Schematic chart of the Next-generation industrial 3D printers project(B) Development of 3D printing systems for sandcasting cores and moldsThe large-scale AM machines and materials for sand mold, which enable us to produce the requiredcomplex molds and unified cores for metal casting were developed by CMET INC..By FY2016, the first large AM machine with an organic binder and a coated sand was developed toachieve the desired build speed of 100000 cc/h and work size (more than 1000 x 1000 x 600 mm).In FY2017, the non-organic binder sand mold for AM system was developed to improve environment of foundry.

Sound cast parts were able to be obtained using this sand. The sand mold AM machine which can use two types of sand partially was developed. The effect of reducing casting defects by the difference of heat capacity in sand mold parts was able to be verified.In FY2018, the largest sand mold AM machine (Fig.20) was developed and achieved the desired buildspeed of 100000 cc/h and work size (1800x1000x750). An example of a sand mold part is shown in Fig.21.Fig.20 Sand mold AM machine developed by CMETFig.21 An example of sand mold1649AcknowledgementThis paper is based on results obtained from a project commissioned by the Ministry of Economy, Tradeand Industry (METI) and the New Energy and Industrial Technology Development Organization (NEDO).References [1] www.meti.go.jp/english/press/2014/0221_02.html. [2] www.trafam.or.jp [3] H.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDEVELOPMENT, PRODUCTION AND POST-PROCESSING OF A TOPOLOGY OPTIMIZED AIRCRAFT BRACKETHelge Klippstein1, Anne Duchting2, Thomas Reiher2, Florian Hengsbach1, Dennis Menge1, Hans-Joachim Schmid11 Direct Manufacturing Research Center and Particle Technology Group, Paderborn University, Germany 2 simufact engineering GmbH, Paderborn, GermanyContact: Helge.Klippstein@dmrc.deAbstractStructural parts for aviation have very high demands on the development and production process. Therefore, the entire process must be considered in order to produce high-quality AM metal parts. In this case study, a conventional part was selected to be optimized for AM.

Therefore, the entire process must be considered in order to produce high-quality AM metal parts. In this case study, a conventional part was selected to be optimized for AM. The process presented includes component selection, design improvement with a novel approach for topology optimization based on the AMendate algorithm as basis of MSC Apex Generative Design, component production on a SLM 250 HL and post-processing including heat treatment and surface smoothing. With the topology optimization a weight reduction of ~60 % could be realized, whereby the stress distribution is more homogeneous. Furthermore, the challenges of support optimization and post-processing have to be addressed, in order to produce competitive parts.IntroductionThe additive manufacturing technology is characterized typically by the multi-layer composition of the components.

Altogether, the research and application area for additive manufacturing is a growing market with often so-called "endless potentials". This can be deduced from the number of publications and patents which have increased significantly since 2014 [1]. At the same time, the number of AM technologies and companies continues to grow. Additive manufacturing is particularly attractive, if the part is highly complex and/or the number of parts per series is relatively small [2]. Applying Design for Function instead of the traditional Design for Manufacturing approach and including part consolidation and integration, increases the complexity of parts and assemblies. At the same time, weight, maintenance and waste reduction can be achieved by using generative manufacturing techniques. In addition, the combination of CAD/CAM and AM will shorten development times [3].

In addition, the combination of CAD/CAM and AM will shorten development times [3]. In aviation, this as well as the weight saving potential or the reduction of tooling and warehouse costs should be taken into considerations by calculating a business case for AM parts [4, 5]. The production of conventional part designs with AM technology contradicts the overriding objectives of AM and aircraft design. Only by leveraging the design freedom of AM technology this process will be truly competitive, raising both performance potentials and cost savings [6]. This is as well reflected in the new approach by Airbus, targeting clean sheet design solutions for the AM implementation [1].Due to the high quality requirements and regulations in the aviation sector, the implementation of AM technology is closely associated with very expensive material and process qualifications.

The1932ASTM F42 and ISO TC 261 committees, namely the most important qualification committees, work on universal guidelines for the AM qualification process [7]. However, up to now there are only a few materials and processes qualified. Today, Boeing and Airbus have already implemented thousands of generative manufacturing parts for civil and military systems. Since ULTEM 9085, PEEK and PEKK are one of the first AM materials suitable for aviation, most of these parts are manufactured using FDM technology. However, the trend towards more metal parts is obvious as several metal AM parts today are qualified and flying. In general there are three approaches for qualification:a) Statistical qualification b) Equivalence-based qualification c) Model based qualification.Whereby, version a) leads to costly empirical tests, the equivalence-based qualification is linked to the statistical qualification and reduces therefore the required amount of tests, e.g.

for a new material on the same machine. The model based qualification is supported by an in detail understanding of the process and related simulations [8]. For example, to certify the GE9X T25 sensor and LEAP fuel nozzle from GE, massive pre-testing was done to guarantee repeatable results. Hereby, the overall process is qualified for this part [1, 9]. However, with progress done in process qualification, the door can be opened for equivalence-based material or part qualification. Furthermore, with improved process simulations, even validated model-based qualification with high accuracy can be realistic in the future. Additionally, the concept of “qualify as you go” might be applied. Hereby, multiple pre-, in- and post-process measurements shall demonstrate that a part will perform as specified [8]. As soon as more AM companies, processes, parts and materials are qualified, the spare part market might become more important.

As soon as more AM companies, processes, parts and materials are qualified, the spare part market might become more important. In future, the spare part could be manufactured on site and therefore reduce warehouse and shipping costs [3, 5, 10, 11]. To reach this bright future scenario, it is important to increase the number of flying AM parts further more. When using AM on existing parts, the biggest challenge is to identify promising candidates. The search for these parts is much more difficult than setting up a new part from scratch, as the business case must cover all change costs, including changes to manufacturing drawings, component lists, possible fastener adjustments, quality checks, maintenance and repair guidelines, etc.

However, within this study an existing aircraft bracket was identified and redesigned for additive manufacturing based on the Laser Beam Melting (LBM) process.Methodology and Original BracketThe bracket identified here is installed less than 100 times a year. It is an assembly of two aluminium brackets milled from solid raw material with a buy-to-fly ratio of approximately 6.25. The boundary box of this part is around 100 x 110 x 120 mm. Based on the part size, cost above average for conventional brackets and small number of parts per year this assembly was chosen for redesign and replacement by an AM part. Related to assembly situation, only the load direction for the attacking forces is known but not the size of those forces. Starting from the load direction, the load capacities and limiting loads are determined by reengineering (Figure 1). The two brackets are preassembled by rivets and fastened to the structure.

Starting from the load direction, the load capacities and limiting loads are determined by reengineering (Figure 1). The two brackets are preassembled by rivets and fastened to the structure. A FEA is executed for each load case with an arbitrary load of 1 N.

Several aviation and AM companies do work hard for the qualification of this material [12]. The part manufacturing is done with an SLM-250HL from SLM Solutions. First a topology optimization (TO) is carried out. This design is further adapted to the LBM-related design guidelines and then manufactured. The parts are reworked and all functional areas are milled as part of the finishing process. Finally, a static test validation of the fixture is carried out on a uniaxial tensile testing machine.AlSi10Mg - Material PropertiesFor the topology optimization and the FEA validation a close look has to be given to the fatigue behavior of AlSi10Mg LBM parts, as still most parts fail due to dynamic loads [13]. Based on the properties of the conventional bracket, dynamic load cases have been calculated and applied to the TO design. A fatigue strength at the ratio R = (cid:653)min/ (cid:653)max = 0.1 is required for the new design.

A fatigue strength at the ratio R = (cid:653)min/ (cid:653)max = 0.1 is required for the new design. LBM aluminium parts are known to have a higher yield and ultimate strength, compared to conventional aluminium parts. The microstructure of LBM “as build” parts is very fine due to the extremely high cooling rates [14, 15]. In contrast, the elongation at break and fatigue strength is in the as build condition mostly lower than the conventionally manufactured counterpart. Several approaches have been made to improve the fatigue resistance of LBM parts. Most promising are surface smoothing and heat treatment. As well, remaining powder moisture has an influence on the porosity and the correlated part density which in turn influences the life time [16]. Obviously, there are multiple additional influencing factors for the part quality. The fatigue information gathered in Table 1 are based on literature values with different treatment methods.

Obviously, there are multiple additional influencing factors for the part quality. The fatigue information gathered in Table 1 are based on literature values with different treatment methods. There are samples with different scanning parameters, layer thicknesses, build platform temperatures, machines or orientations included to the table.

In every case only the weakest sample set is displayed per treatment method.1934(1)(2)(3)Table 1 AlSi10Mg fatigue literature values of the weakest tested samples for various post treatment methods *“T6 heat treatment: solution heat treated for 1 h at 520 °C followed by water quenching to room temperature and then aged for 6 h at 160 °C” [17] ** no full Wöhler curve run *** High Isostatic Pressure (HIP) treatment for 2 hours at 180 MPa **** “~” Values are read from graphsSource Load Ratio R = 0.1[15][17]R = 0.1[18] [19][20]R = -1 R = -1 rotating beam mode R = -1MachineTrumpf TrumaForm LF130Renishaw AM250 SLMEOS M400 unknownEOSINT M-280Cycles WeakestPost treatment(cid:653)max~5.5x106 5x107 5x1077.6x106 3x107 3x1072x107 1x107~90 MPa machined ~100 MPa machined ~140 MPa T6* heat treated and machined 63 MPa as build 94 MPa T6* heat treated 126 MPa T6* heat treated and machined 140 MPa machined as build ~50 MPa1x107 3x 107~90 MPa polished50 MPa machined and polished1x 107~75 MPa+ HIP*** 500°C stress relieved (300°C / 2h) machined and polished[21][22]**R = -1 rotating beam mode R=0.1EOSINT M-270EOSINT M-2801x 107 5x1061x105~125 MPa machined or polished as build 97±7 MPa178 MPa** stress relieved (300°C / 2h)machined3 x104222 MPa** stress relieved (300°C / 2h)machinedWith the information of Table 1 the maximal stress for the topology optimization and FEA validation can be deducted.

EOS GmbH and SLM Solutions AG suggest a heat treatment for two hours at 300°C, as the typical microstructure of LBM parts is already close to the one obtained after T6 heat treatment [21, 23]. In contrast, Uzan et al. [20] have experienced a life time reduction by the stress relief procedure whereby large pores within the fracture surface were observed. However, the life cycle points checked by Tang et al. do indicate a much better performance of parts annealed by stress relief [22]. The large variety of obtained mechanical properties substantiate that the effects of e.g. machine- and exposure setting, powder properties and testing procedure have significant influences on the results. Most of the data presented in the table are related to the stress ratio of R = -1, whereby the bracket under consideration here is loaded by the ratio R = 0.1, which in general allows higher maximal loads than at R = -1 ratio. Within the post processes a stress relief heat treatment is planned.

Within the post processes a stress relief heat treatment is planned. Even surface smoothing operations will be investigated. Thus, for simulation of the bracket a fatigue1935strength of 80 MPa has been assumed, which appears to be well in the range of the literature values and covers uncertainties of process parameters with a factor of 1.17 for just heat treated [17] and 1.25 for machined samples from [15]. The used material properties are related to the powder, machine and exposure settings.

Combining the dark green and dark red bracket much more three-dimensional optimization potential is generated, as well as the benefits of part reduction. The bracket has five bolt connections to the airframe, as marked in yellow in Figure 2.Figure 2: Left graphic: Conventional bracket composition with main bracket (dark green) and additional bracket (dark red). On the right: Design space (green) and non-design space (red) for topology optimizationFor the design optimization there is a general principle used that is called “topology optimization” (TO). Topology optimization is an automated extension for FEM based structural analysis. With this iterative procedure a technical component can be optimized by changing its geometry regarding different characteristic values. Most of the commercially available TO-software offers optimization goals such as mass reduction at a given stiffness or a stiffness maximization at a permitted mass.

Most of the commercially available TO-software offers optimization goals such as mass reduction at a given stiffness or a stiffness maximization at a permitted mass. At the same time, the permitted strength must not be exceeded to ensure the structural integrity. For most TO-processes, an additional FE-analysis is needed to validate the occurring stresses of the final design.1936The process of topology optimization starts with a model setup, according to the classical FE analysis including loads and external constraints on all interface areas. To run a proper topology overall geometry is divided in a “design space” where material is allowed but not mandatory, and the “non-design space”, where material is mandatory as for interfaces such as bolts or bearing seats. To enable the optimization algorithm to generate best designs, it needs to have the most possible design freedom, what is achieved with a design space that is as large as possible.

To enable the optimization algorithm to generate best designs, it needs to have the most possible design freedom, what is achieved with a design space that is as large as possible. It can be a simple but massive geometry or an arbitrary complex one as in complex assemblies, which defines the maximum allowed material for the component. By solving the FE calculation, stress and strain can be determined for each of the elements located in this design space. The optimization algorithm now decides which of the elements can be removed without exceeding given restrictions such as displacement or stress maxima. Subsequently, a new analysis is performed with this reduced amount of material, the element stresses are interpreted, and further material is removed or reattached. Performing this iterative procedure, the optimization algorithm approaches an optimal component shape step by step.

Performing this iterative procedure, the optimization algorithm approaches an optimal component shape step by step. For this part the design space is created considering the adjacent components and is designed as generously as possible. It should also be noted that the nondesign spaces must be accessible for assembling and therefore material must be left out. The result can be seen in Figure 2 with the design space marked green and non-design space marked red. The software solution “AMendate” was used for the optimization of the bracket. AMendate is a topology optimization software that has been developed by a young start-up of former DMRC- employees from Paderborn University and is now part of the commercially available software solution “MSC Apex Generative Design 2019”. The main difference of this “generative design” tool towards most known topology optimization tools is the automated interpretation and definition of a final geometry during the optimization.

The main difference of this “generative design” tool towards most known topology optimization tools is the automated interpretation and definition of a final geometry during the optimization. This enables the use of a clearly defined stress goal instead of the beforehand mentioned ones: mass reduction or stiffness. As for other optimization tools, during the optimization the material is reduced based on the FE results. Though, in each iteration a specific design with a geometrically and mechanically useful surface is generated. Thus, the user does not need to interpret the results and do manual redesign in a CAD-system. Each design is analyzed by the algorithm regarding the occurring stresses in detail. This is possible due to a completely new approach that is not based on the popular SIMP-algorithm (solid isostatic material with penalization) which gives a varying density to each element, but on a so called hard- kill algorithm.

In these algorithms the elements considered to be unimportant for the structure with low stress or strain are fully deleted and not considered in further iterations. This hard decision enables a generation of a high quality part surface and thus determination of a clearly distinct structure [28]. The software allows to influence the generated design in terms of complexity like a more “filigree” design with many small struts and thin shells or with a more “massive” design with less but thicker elements. Three versions have been calculated as shown in Figure 3. Hereby, the violet (rough) and orange (normal) bracket do variate in weight about 0.1 g, whereby the blue (filigree) bracket weights 2.7 g more than the orange one. All brackets have similar maximum stress levels and do not violate the 80 MPa limit. The stiffness of the orange one is the highest. An additional criteria for the design selection is the amount of attachment points.

The stiffness of the orange one is the highest. An additional criteria for the design selection is the amount of attachment points. The violet (rough) design has only three bolt connections to the airframe. For safety considerations, it can be expected, that additional fastening connections and therefore struts will lead to a more good-natured failure pattern. Hence, the normal setting is used for further development, it is based on stress and stiffness very good, has multiple fastening points and as well the structure is not too filigree.1937Compared to bulky parts, thin structures do show much weaker mechanical properties due to an increased ratio of contour to volume exposure and therefore increased porosity [14, 24]. These imperfections should be taken into account when evaluating the results of topology optimization. In particular, the very filigree struts cannot withstand the calculated loads of the FEA.

These imperfections should be taken into account when evaluating the results of topology optimization. In particular, the very filigree struts cannot withstand the calculated loads of the FEA. Therefore, an offset is given to the overall design and all functional areas are thickened to allow subtractive post-processing. The areas subject to higher loads are additionally reinforced. The manual rework is done together with the support optimization of the bracket.Figure 3: Three different TO results from left to right with more filigree structures. Based on the filigree settings three (rough result – left bracket) up to five (very filigree result – right bracket) attachment points are used. The final design with is developed further is the orange one with four attachment point.Support OptimizationFor the LBM manufacturing process a supporting structure is required.

The final design with is developed further is the orange one with four attachment point.Support OptimizationFor the LBM manufacturing process a supporting structure is required. Based on the AM design guidelines, the build orientation should be specified at the beginning of the design process, since it has several influences on the component properties [25]. For topology optimization without a support optimization algorithm, however, the build orientation must be checked first. The software Magics from Materialise and Meshmixer from Autodesk are used to optimize the build orientation. Within this software, the STL file and orientation can be analyzed and optimized for various purposes. These algorithms are based on the triangle orientation of each surface element in the STL file. This allows a threshold overhang angle to be defined related to the process and the material. All triangles with an orientation smaller than this angle should be checked for support requirements.

This allows a threshold overhang angle to be defined related to the process and the material. All triangles with an orientation smaller than this angle should be checked for support requirements. The objective of the optimization used here is to minimize the summed up projected area on the X-Y plan. For post-processing, the inner support structure, which grows from part to part, is particularly critical as it is difficult to remove and leaves double marks on the component surface, one on the underside and one on the top. In the LBM process, the purpose of the support is not only to keep the material in place, it is also required for heat dissipation. With aluminium, the cone support in particular fulfils this purpose, as the contact surface of the line or block support with tooth is very small. Experience shows stable structures with good surface properties up to an overhang angle of 35°.

Experience shows stable structures with good surface properties up to an overhang angle of 35°. In the literature even smaller angles can be built, but the surface roughness increases significantly [26]. Newer LBM machine generations do support special overhang parameter settings whereby the trend is even going to much smaller angles with different exposure strategies and laser controls, e.g. pulsed laser [27].1938The purpose of the following procedure is to adapt the part design, in order to reduce the overall support area. Particularly, the struts supported internally should be removed. Small holes and voids are filled up. The part is sculpted to reduce overhangs to a limit of 35° which is shown in Figure 4. Within this procedure the orientation angle is checked again and fine part orientation adjustments based on the small modifications are done.

Within this procedure the orientation angle is checked again and fine part orientation adjustments based on the small modifications are done. The comparision of the original TO version and the support addapted one shows significantly smaller overhang areas and therefore less required support. Areas with two struts merging together, called bridging, do not need to be supported as the part itself supports the structure efficiently. Aliminium is not very sensitive for residual stress deformation during the build process, still additional cone supports has been added to the functional areas of the bracket to reduce deformation due to residual stress.

Test build jobs with both designs have shown, that the postprocessing of the support adapted design is more than 10 times faster, as the standard TO design and takes around 2 minutes for the rough support removal and 15 minutes for mechanical smoothing of the attachment areas.Figure 4 Optimized build orientation with reduced support area and less numbers of merging struts; Overhang angle of 35° marked in red. Top row displays the TO output brackets, bottom row shows the adapted design. Left: close ups of examples for the modifications: filled up voids and sculptured struts to avoid part to part support. Middle: Bottom View, support surfaces in red, right: bottom view of the applied support structureFinite Element AnalysisFinally, the results of the topology optimization and the part adaptions for quickly removable support structure, need to be validated. Based on a finite element analysis, the part is tested for static and fatigue behavior.

Based on a finite element analysis, the part is tested for static and fatigue behavior. The software Solidworks 2018 from Dassault Systèmes was used for the validation.1939Load 1Load 1Load 245% 42% 38% 34% 30% 26% 23% 19% 15% 11% 8% 4% 0%Load 270% 64% 58% 52% 46% 40% 35% 29% 23% 17% 12% 6%Figure 5 TO-Bracket validation, left Von Mises stress (normalized to maximal stress of the original bracket) right: deformation (normalized to maximal deformation of the original bracket)By comparing the FEA results of the TO bracket to the FEA of the conventional bracket, it becomes clear, that the stiffness of the part has increased. The new design can hold three times higher loads of the first load case and 2.1 times higher loads in the second load direction. Therefore, load two is the limiting load, as the initiated stresses within the bracket are higher. Hence, load two is used as the validation load for the bracket tested on a uniaxial tensile test machine.

Hence, load two is used as the validation load for the bracket tested on a uniaxial tensile test machine. Combining both loads, the maximal stress of the bracket reach 114 MPa for the ultimate load and thus 75 MPa for the limit load. Applying disturbance loads by vary the load direction of both loads about 10° shows a maximal stress increase of ~25% to 96 MPa. After further thickening of the highly loaded struts a rerunning of the analysis reveals that the disturbance load impact is reduced to 5% stress increase. The bracket is therefore designed close to the limit and will withstand the limiting fatigue stress level of 80MPa. However, in order to master the ultimate load in cyclic loading, the bracket should be finished furthermore.

However, in order to master the ultimate load in cyclic loading, the bracket should be finished furthermore. Another option would be to further inflate the highly loaded struts.Manufacturing, Post Processing and FinishingThe brackets are manufactured on a SLM 250-HL from SLM Solutions which is running on build processor 2.2 and cannot apply overhang parameter settings. As the powder moisture does have a significant influence on the part quality, the powder was dried prior to the process. With dry powder the hydrogen porosity can be reduced significantly as well as a better part surface can be obtained due to the reduction of weld spatter [16]. 35 cycles of vacuum drying and argon rinsing can dry around 5 kg of aluminium powder within 35 minutes. The machine runs for 26 hours with a job height of 110 mm to manufacture eight parts on one build platform. These work cycle times can be reduced on modern multi-laser machines to improve the business case.

These work cycle times can be reduced on modern multi-laser machines to improve the business case. To reduce tension and potential deformation, as well as to obtain a higher ductility of the brackets, the build brackets mounted on the build platform are stress relief annealed. After removing the support structure and smoothing the attachment marks manually, all functional surfaces, as displayed in Figure 6, are milled. All brackets are blasted with corund, followed by glass bleads.1940Figure 6 Functional surface of the bracket in red and overview of the brackets during unpacking of the build jobStatic Test ValidationFor the static test of the bracket, it is loaded in the direction of load case two which indicates higher stresses within the bracket. The specimen is fixed on a massive steel block at the bottom of the tensile test machine. As displayed in Figure 7, the load is applied via attachment point B in an angle of 30°.

The specimen is fixed on a massive steel block at the bottom of the tensile test machine. As displayed in Figure 7, the load is applied via attachment point B in an angle of 30°. The attachment point A for load one is bolted together as a part of the stresses applied by load two will be guided through this connection, which has a bolt connection even in the real assembly. All fixing points from the bracket to the rig as well as the load attachments are prepared with engineering fit H7. Especially in the beginning of the load curve the gradient is not linear (Figure 7). It is expected, that not all struts are loaded equally, thus the final stiffness of the bracket is reached first at around 80% load. The required ultimate load of the conventional (original) bracket is withstood by the new design without fail. The component has passed the test for the calculated limit and ultimate load intact, as well. At 225% load the limits of the test set up are reached.

The component has passed the test for the calculated limit and ultimate load intact, as well. At 225% load the limits of the test set up are reached. Thus, the bracket could not be run in a destructive test.l250%a n o i t n e v n o C e h t)%200%150%Calculated ultimate load of the tested TO bracket(f o o i t a r d a o L100%50%Ultimate load of the conventional bracket0%00,250,51 Traverse movement (mm)0,751,251,51,75Figure 7 Static validation of the support optimized TO-bracket, Left: the bracket in assembled on the test rig for load 2 attacking in 30° on attachment point B. Load 1 attachment point (A) is bolted together to simulate the additional stiffening connection. Right: Topology optimized bracket test diagram.

Load 1 attachment point (A) is bolted together to simulate the additional stiffening connection. Right: Topology optimized bracket test diagram. Test stop without failure at 225 % of the load related to the original (conventional) bracket.19412In combination with the engineering fit connections and the assumption, that the test rig made of steel and fixed by steel 8.8 screws is much stiffer than the aluminium bracket, it can be expected that the traverse movement is equal to the elongation of the bracket in load direction. Here it is essential, that the elongation for the 100% load of 0.65 mm is much higher than calculated by the FEA. Whereby, with 59 GPa the used E-Modulus for the FEA material model is already conservative. Additionally a fatigue test of the bracket design has been run at 20 Hz with a tension load ratio of R = 0.1 for load case two. Hereby the load pattern follows a sinus curve.

Additionally a fatigue test of the bracket design has been run at 20 Hz with a tension load ratio of R = 0.1 for load case two. Hereby the load pattern follows a sinus curve. The test was performed on a Zwick/Roell HC10 servohydraulic testing machines, whereby the bracket is bolted to the same test rig, as used for the static test. After ~4,670,000 cycle loads the test was stopped due to a violation of the lower stress limit. By restarting the test some massive stress limit violations led to a total fail of the part short after the restart of the machine. SEM images has been taken to analyze the failure pattern as presented in Figure 8. Here the crack initiation, propagation as well as resting lines and the final fracture zone can be seen. Very prominent are the amount of pores and unmolten particles, exemplary marked in violet at the bottom left image of Figure 8. It was found, that the used SLM machine had a defect laser, which could not deliver the set laser power.

It was found, that the used SLM machine had a defect laser, which could not deliver the set laser power. Hence the part validation has to be repeated. However, even with those drawbacks the bracket has withstood close to 4.7 million cycles as well as the static stress tests.Crack startFigure 8 Bracket after dynamic tests. Top left: crack line through the bracket. Top right: SEM image of the fracture zone. Bottom left: crack initiation point. Bottom middle: rest lines and Bottom right: final fracture region1942Conclusion and OutlookIt has been shown, that performance improvements can be achieved by topology optimization and additive manufacturing. The support optimization for the LBM process is hereby still in focus, which leads to manual design adaptations. However, the company AMendate is working on an optimization algorithm for their software, to generate support optimized TO results. This in turn will reduce the required design effort after the TO dramatically.

This in turn will reduce the required design effort after the TO dramatically. Furthermore, the individualities of the LBM process need to be considered for every FEA analysis, as these structures often show imperfections, which are related to a complex process. Powder drying helps already to improve the process stability and part quality. Still differences between simulations and real build part tests can be seen, e.g. in the predicted fracture region. Hereby the impact of the lately detected laser defect could be a reason for this and at least for the high porosity within the parts. Hence, at least for machines without a melt pool monitoring system additional test samples are recommended for every build job with critical parts. The work done for this bracket shows the potential of AM.

The work done for this bracket shows the potential of AM. As soon as the material is qualified for civil aviation applications, multiple brackets like this, could be replaced as weight savings of around 60% can be achieved additionally with improved performance. Statically test have shown, that the material model used for the FEA simulation seems to be conservative. The dynamic testing of the bracket is a very important step towards AM structural parts in aviation and need to be repeated for this bracket. Only if the high cycle fatigue demands are fulfilled, this part will fly.AcknowledgementThe authors would like to thank the German Government, the Federal Ministry for Economic Affairs and Energy (BMWi), all involved industry partners and the Paderborn University for the support within the project VERONIKA: 20K1513D.ReferencesReferences [1] T. Wohlers, R. I. Campbell, and O.

E RECONFIGURATION AND RESTORATION USING LASER-BASED DEPOSITIONT.W. Skszek and M. T. J. LowneyAbstractPOM Company, Inc., located in Plymouth, Mich., has successfully commercialized the laser-based, free form fabrication process known as DMD,™ or Direct Metal Deposition. Since the company incorporated in 1998, POM has been directed towards the commercialization of the DMD process equipment, and the demonstration of value-added, cost-effective applications specific to the fabrication, repair and reconfiguration of plastic injection molds and die cast tooling. POM currently operates as a service bureau and original equipment manufacturer (OEM), providing engineering analysis, design and fabrication services, as well as the sale of DMD process equipment.

POM currently operates as a service bureau and original equipment manufacturer (OEM), providing engineering analysis, design and fabrication services, as well as the sale of DMD process equipment. An overview of the POM closed-loop, laser-based DMD process, available services and several case study applications are presented.DMD Process OverviewThe closed-loop DMD process includes a CO2 laser and CNC-controlled overhead gantryto manipulate the focusing optics and powder feeder nozzle according to CAM tool path data associated with a solid-model CAD geometry.

In contrast to the selective laser sintering (SLS) process, which utilizes a laser to sinter a variety of powder-based plastic materials to form 3-D shapes and prototype molds, the DMD™ process uses pure metal alloy powder to form 3-D metallic shapes.Metallic powder, typically tool steel alloys (P20, S7, H13, H13+, 420SS or 420SS), copper alloys (CuCr, CuNi, CuSn or CuZn) or OFHC pure copper, is transported to the nozzle by an inert gas (Ar or ArHe) at a pre-defined rate from one of three on-board powder feeders. The metallic powder is added to a melt pool established by a focused laser beam, concentric to the nozzle. POM’s patented closed-loop feedback control system maintains the size and shape of the melt pool, providing dimensional control and a very high quality, near-net shape deposition.Moving the laser heat source and nozzle at a controlled rate according to the pre-definedCNC parameters and CAM path geometry results in a defined cooling rate (103 -> 106 degrees/sec).

This provides uniform hardness, microstructure and isotropic material properties throughout the volume of deposition. The CNC tool paths and laser process parameters are tailored to achieve the desired cooling rate and as-deposited microstructure. Depositing tool steel alloys using the DMD™ process typically results in an “as-quenched”, fully hardened, martensitic microstructure with very fine and uniform distribution of intermetallic carbide particles. This provides improved mechanical properties (toughness and yield strength) as compared to those of commercially available wrought alloys. Akin to that of laser welding, the heat affected zone (HAZ) associated with the additive metal deposition is very shallow (typically2190.015”), and void of the typical “sink” type defects associated with conventional high- temperature welding processes.The as-deposited material is typically tempered, using conventional thermal heat treatpractice or a defocused laser beam.

Laser tempering provides the benefit of “color matching” the microstructure of the deposition to that of the substrate material. It also reduces processing time and costs associated with disassembly, pre-heat and post-heat treatment processes.From an engineering perspective, a customer typically transmits a 3-D solid model of themold or die to the POM ftp website (www.pom.net) for feasibility evaluation and quoting. A solid model associated with the desired deposition volume is created and sliced into layers, which coincide with the height or thickness of layered deposition process (typically 0.005” -> 0.015” thick). CNC tool paths, identical to those used for CNC machining, are then generated for each layer using traditional CAM software (SDRC-Ideas). The CAM output is then post- processed, using POM proprietary software. This software basically inverts the CAM file and imbeds laser and powder commands specific to the application requirements.

The CAM output is then post- processed, using POM proprietary software. This software basically inverts the CAM file and imbeds laser and powder commands specific to the application requirements. The post- processed file is then downloaded to the DMD™ machine for processing.A 3-D shape is fabricated on a layer-by-layer basis. It is identical to that of the CAD model used to generate the tool paths, by depositing “roads” which are approximately 0.040” wide. The deposition rates, specific to each material, range between 0.50 in3/hr and 0.75 in3/hr. The relatively low deposition rate, and low heat input (Btu/hr) associated with the process, provide conditions that are ideally suited for additive metal repairs and fabrication of composite material molds and dies.The case study applications presented illustrate the commercial viability of the DMD™process.

These case studies include the deposition several tool steel alloys associated with die restoration, mold reconfiguration, repair of plastic injection molds and die cast tooling. Each application demonstrates how this technology can reduce tooling costs and improve tooling lead times.Case Study ApplicationsDie Cast Tool Restoration – The functional use of die cast tooling (die life) is commonly limited by the initiation and growth of cracks due to “heat checking.” Heat checking results in brittle failure or the entrance of cooling water into the die cavity. This type of failure is commonly known to result from thermal fatigue imposed by high thermal gradients, thermal cycling and the reduction in mechanical properties (impact strength and yield strength) at elevated operating temperatures.

It is not surprising then that heat checking type failures are often observed in areas of the die that are difficult to cool (elevated die temperature), subjected to high metal velocities (elevated die temperature), and/or adjacent to a water line (thermal gradient).A significant portion of the cost associated with a die cast part is related to tooling costand die life. Typically, the cost of the tooling is amortized over the anticipated life of the die and220added to the piece part price. Due to the lack of an acceptable restoration process, dies are frequently comprised of inserts, which are replaced throughout the life of the die. In many cases, the part design or die design does not permit use of replaceable inserts as failure takes place in the die block, requiring purchase of new tooling.In most cases, the heat checked or eroded areas of a tool are confined to localized areas, which are difficult to cool and subjected to high thermal gradients.

The DMD™ process offers significant cost savings potential associated with restoration of heat checked and eroded die surfaces. DMD™ affords this opportunity because it can restore existing tooling, rather than requiring tooling to be replaced…and, at a fraction of the cost.Case Study #1The ingate and runner areas associated with four severely heat checked H13 die cast inserts were restored by POM using the DMD™ process. Further production use and related thermal cycling would have resulted in replacement of the inserts due to catastrophic failure or propagation of the heat check induced cracks into the water line.The DMD™ laser-based restoration solution included:Removal of approximately 4 in3 of thermally fatigued, heat checked material from the runner area of each of the four inserts.

This was accomplished using conventional CNC machining.Deposition of H13 tool steel, restoring the heat checked surfaces of two insertsDeposition of POM proprietary H13+ alloy, which possesses improved mechanical properties at elevated temperatures, was deposited in the runner area of two inserts.Thermal tempering.CNC machining of the inserts per the original CAD geometry.The four-cavity mold is currently being used in production. The results of the on-going production trial, which commenced in July 2000, will provide valuable insight on the use of the DMD™ process to restore heat checked die surface. It will also serve as a comparative evaluation of two-alloy compositions, relative to the potential increase in die life. The benefits this company has realized already include a significant reduction in tooling costs realized through increased die life through restoration vs.

The benefits this company has realized already include a significant reduction in tooling costs realized through increased die life through restoration vs. replacement of die cast tooling.221Figure 1 – Pre-machined, heat checkedFigure 2 – DMD-restored insertrunner areaReconfigurable Tooling – The product development cycle associated with plasticinjection molded parts historically includes the procurement and fabrication of “soft tools” to produce prototype parts. This is followed by the design and build of production tools, once the product design is finalized. Until recently, “soft tooling” was comprised of aluminum, which can be CNC machined at high feed rates and easily reworked to accommodate engineering changes during the product development cycle.In recent years, due to the advances in high-speed CNC machining, an increased percentage of soft tools are fabricated using free-machining, low-carbon steel.

Fabrication of low-carbon steel prototype tooling maintains the low cost and short lead-time benefits associated with aluminum tooling. Simultaneously, it provides the molder with inset associated with the thermal and filling aspects of the molding process due to the similar thermal conductivity of steel alloys. Although significant improvements have been made in recent years, the overall cost and lead-time associated with two tooling development programs has not been addressed.The DMD™ laser-based deposition process provides the capability to fabricate composite material molds comprised of low carbon steel substrate and tool steel mold surfaces. This combination offers the low cost and compressed lead-time benefits of low-alloy steel tooling, while providing hardened tool steel mold surfaces required for production use.

This combination offers the low cost and compressed lead-time benefits of low-alloy steel tooling, while providing hardened tool steel mold surfaces required for production use. In essence, the reconfigured prototype tool becomes the production tool…at a fraction of the cost and lead-time required to design and build a prototype tool and a production tool.222Case Study #2POM recently reconfigured an obsolete two-cavity, low carbon steel prototype tool usedfor production of automotive body side moldings.

The tool reconfiguration process included:Offsetting the CAD geometry associated with the new part geometry by 0.060”.Overlaying the existing mold geometry with that of the new part.Generation of tool paths.CNC machining of the obsolete mold surface.Laser-based deposition of 0.090” thick layer of P20 tool steel over the 1025 steel substrate.Thermal tempering of the P20 deposition.Re-machining and benching of the styling surface.Prototype parts were submitted in 35 days, at a fraction of the cost and lead-time it takes to design and build a new prototype tool.

The typical lead-time associated with the construction of conventional prototype tooling is 12 to 14 weeks.Figure #3 – Pre-machined 1025 LCS (laser clad surface) of body side molding prototype tool223Figure #4 0.090” P20 tool steel styling surfaceAdditive Metal Deposition – All too often, machining errors or last-minute engineering changes jeopardize on-time delivery of the tool and potentially impact the introduction date of a new product. Frequently, new tooling is “welded” using conventional processes and the part quality standard is compromised until a replacement cavity can be manufactured.The high-temperatures associated with conventional welding processes impose variablecooling rates throughout the relatively large heat affected zone (HAZ). This results in composite microstructure comprised of martensite, bainite and retained austenite.

This results in composite microstructure comprised of martensite, bainite and retained austenite. In addition, due to the lack of adequate shielding and the presence of deoxidizing elements (Al, Ti & Si) in the weld rod, the weld deposition often includes a high percentage of oxide inclusions and intermetallic compounds. This reduces mechanical properties, hard spots and porosity. In an effort to minimize weld stresses developed as a result of solidification shrinkage, and to normalize the variation in cooling rate, conventional tool steel welding procedures typically specify pre-heat and post-heat treatment of the mold or die. “Color match” weld procedures specify post-weld solution heat treatment and tempering.

“Color match” weld procedures specify post-weld solution heat treatment and tempering. These procedures often induce dimensional distortion, cracking and “sink” type defects at the weld interface, resulting from the transformation of the composite microstructure to martensite during post-weld heat treatment.224The DMD™ process is ideally suited for additive metal deposition applications because ituses pure alloy powder (void of deoxidizing elements) and low heat input specific to the laser- based deposition process. The as-deposited martensitic structure, shallow heat-affected zone and laser tempering provides a deposition that is identical to that of the parent material, and does not require preheat and post weld heat treatment.Case Study #3POM recently restored an S7 tool steel production mold insert used in the production of an automotive thermostat manifold.

The 1.8 in3 3-D section, which had broken away from the solid during production use, was redeposited using the DMD™ process. The tool reconfiguration process included:Pre-DMD machining of the irregular surface.Generation of tool paths from the customer-supplied CAD geometry.S-7 tool steel deposition (restoration) of the 3-D shape.Solution heat treatment & tempering.EDM finishing of the near-net shape deposition.The restored mold was returned to the customer in three days for final finishing and wasback in production in a total of five days. All was accomplished at a fraction of the cost and lead-time associated with the build of a new tool insert. Due to the lack of a back-up tool, which is not uncommon in the molding industry, the customer ran one of its two-cavity molds, 24 hours/day to supply parts to the customer during five-day tool repair process.

The typical lead- time associated with the construction of replacement tooling at premium rates would have been eight weeks.Figure #5 - S7 thermostat mold Figure #6 - Restored S7 thermostat mold with a broken core (see lower right)(see rebuilt core in lower right)225Case Study #4A late engineering change reduced a wall stock associated with an electrical connector housing. This created the need to add 0.035” of H13 tool steel to the surface 10 new hardened mold inserts (Rc 48).The tool reconfiguration process included:Fabrication of copper masks to prevent the possibility of tool damage due to stray laser reflections.Generation of tool paths from the customer-supplied CAD geometry.H13 tool steel additive metal deposition to the surface of the mold.Laser tempering.EDM finishing of the near-net shape deposition .The stock added to the finished and hardened production inserts provided significant costsavings to the customer.

It also drastically reduced the time required to remanufacture 10 new inserts to accommodate the engineering change.###226.

gital Data Processing Strategies for Large Area Maskless Photopolymerization Anirudh Rudraraju, Suman Das, Georgia Institute of TechnologyAbstract:Large Area Maskless Photopolymerization (LAMP) utilizes scanning spatial light modulators that require layer slice data in the form of high‐resolution bitmaps. Three different strategies have been implemented to fill this need. First, bitmaps were generated by direct slicing of CAD models using Spatial Technology’s ACIS kernel. Second, bitmaps were generated from STL files through ray‐tracing. Finally, an approach involving reconstruction of topological information from STL files for efficient slicing and image generation is being developed. This paper gives a brief description and implementation details of each of these strategies as well as data compression techniques being pursued by the authors.

This paper gives a brief description and implementation details of each of these strategies as well as data compression techniques being pursued by the authors. This work is sponsored by DARPA grant HR0011‐08‐1‐0075.Introduction:Large Area Maskless Photopolymerization (LAMP) technique is an integral rapid manufacturing process which uses a spatial light modulator to expose whole areas in a single flash. The spatial light modulator in our case is a Digital Micro-mirror Device (DMD) chip developed by Texas Instruments. It is essentially a small chip with millions of tiny mirrors that can be turned on and off by feeding an 'exposure bitmap'. Light from the UV-source is shone onto the DMD chip and the various mirrors turn on and off accordingly to project the input image on to the photocurable resin surface.

Light from the UV-source is shone onto the DMD chip and the various mirrors turn on and off accordingly to project the input image on to the photocurable resin surface. The head, consisting of the light source and the DMD chip, raster scans the entire exposure region which gives much higher speeds than conventional prototyping processes where typically a laser beam is used to raster scan the exposure region line by line. Also, because of the high scanning speeds extremely large areas (of the order 100 square inches) can be exposed without compromising on feature resolution. The feature resolution achievable in this process is only limited by the size of the tiny mirrors on the DMD chip (16-17 microns). Thus using this technique, macro sized parts with extremely high resolution features can be obtained with good process speeds.The DMD chip takes high resolution bitmap images as input for exposure. Consequently all the CAD slice data needs to be output in this format.

Consequently all the CAD slice data needs to be output in this format. Most commercial slicing programs output slice data in the form of contours or scan patterns specific to the prototyping process they are designed for. Moreover, they work mostly only on STL files which are just an approximation of the original CAD models thus compromising on feature accuracy and resolution. Also, these commercially available programs do not offer much flexibility to modify the output data to suit the needs of the LAMP process. Hence new data processing algorithms that can efficiently handle both native CAD files and STL files for generating slice data specific to this technique need to be implemented.Also, owing to the high feature resolutions (16-17 microns), slice images need to be generated at very high resolutions (1500 dpi). This can generate enormous amounts of data.

This can generate enormous amounts of data. Each uncompressed slice image is typically of the order of several megabytes, and at this scale theslice data for an entire part can run into several terabytes. Having data on this order of magnitude poses severe storage issues and also hinders fast data transfer between the controlling computer and the prototyping machine. Thus, efficient data compressions schemes and innovative data transfer techniques also need to be investigated.In this paper, the various slicing strategies, both direct slicing and STL file slicing, investigated so far have been presented. The basic approach and the brief implementation details of each of these approaches are described in the sections that follow.Direct Slicing:In the Rapid prototyping industry 3D CAD models are first tessellated and converted into an intermediate file called the STL file format. This format has become the defacto industry standard for slice data generation.

This format has become the defacto industry standard for slice data generation. An STL file just contains a list of all the facets in the model in no particular order. Since it is a mere approximation of the original CAD part, many inaccuracies occur in the tessellated model and hence unsuitable for high resolution builds. Through LAMP, we intend to achieve feature sizes of the order of 16-17 microns and hence the more preferred approach would be to directly slice the original CAD models without the intermediate meshed approximations. Consequently, a direct slicing approach has been investigated. There has been some work in this area of direct slicing in the past. Guduri et al. proposed a direct slicing method for slicing a constructive solid geometry (CSG) representation of a part [1]. Vuyyuru et al. directly sliced solid models built by the SDRC’s I-DEAS and segmented NURBS-based (non- uniform rational B-spline) contour curves [2]. Chen et al. directly sliced PowerSHAPE Models [3].

directly sliced solid models built by the SDRC’s I-DEAS and segmented NURBS-based (non- uniform rational B-spline) contour curves [2]. Chen et al. directly sliced PowerSHAPE Models [3]. Cao et al. proposed a method for directly slicing AutoCAD models [4]. In this paper, we present an approach for directly slicing CAD models using the ACIS kernel.I) Using ACIS:The ACIS kernel is a commercially available 'C++' CAD library marketed by Spatial systems, a subsidiary of Dassault Systemes. It offers robust APIs and function calls for most of the basic CAD operations. These APIs have been integrated into the slicing software to produce CAD slices. The produced CAD slices were then 'rasterized' to obtain the bitmaps used for exposure. The basic algorithm is outlined in Fig. 1.The original CAD part that needs to be sliced is first loaded into the Algorithm using ACIS's load functions.

The basic algorithm is outlined in Fig. 1.The original CAD part that needs to be sliced is first loaded into the Algorithm using ACIS's load functions. ACIS libraries can only work with the "SAT" file format and hence CAD files in other formats need to be converted into the SAT format either by using commercial CAD softwares or by using ACIS's inbuilt file format translation functions. During the translation, numerical or topological inaccuracies could creep into to the part. In severe cases, error checking and correction schemes need to be implemented. Once the part has been loaded, its bounding box is computed to get an estimate of the size of the bitmaps that would be generated. A slicing plane is then created and intersected with the part using ACIS's Boolean APIs to get an intersection wire. Once the intersection wire is obtained, it is "rasterized" to obtain the bitmaps.

A slicing plane is then created and intersected with the part using ACIS's Boolean APIs to get an intersection wire. Once the intersection wire is obtained, it is "rasterized" to obtain the bitmaps. This essentially involves shooting rays for each row of pixels in the image and computing the intersection points with the intersection wire. Pixel values are then filled with alternating white and black segments in between each of these intersection points as shown in Fig. 1 (e). Thebitmap images obtained are then compressed using CCITT fax4 compression scheme which compresses the data by three orders of magnitude without any loss. (CCITT fax4 is an industry standard lossless compression scheme for efficiently compressing 1-bit TIFF images). These bitmaps are then fed to the DMD chip for exposure.Fig. 1: Direct slicing algorithm: (a) Original CAD part loaded into the program. (b) Bounding box computed. (c) Slicing plane created.

These bitmaps are then fed to the DMD chip for exposure.Fig. 1: Direct slicing algorithm: (a) Original CAD part loaded into the program. (b) Bounding box computed. (c) Slicing plane created. (d) Intersection between the part and the slicing plane computed. (e) Intersection wire rasterized to create bitmap images. (f) Slice bitmaps obtained and compressed in CCITT FAX4.STL file slicing:Since STL files are the defacto industry standard and the direct slicing approach using ACIS cannot handle these files, several alternate methods have been investigated. The implementation details of each of these approaches are given in the following sections.II) Using POVRAY:POVRAY is an open source ray tracing software that is routinely used in the computer graphics industry to render high resolution photo realistic images. The schematic in Fig.

The schematic in Fig. 2 shows the principle of ray tracing:A small routine was used to convert an STL file into a POVRAY recognizable mesh A small routine was used to convert an STL file into a POVRAY recognizable mesh A small routine was used to convert an STL file into a POVRAY recognizable mesh object. The object is then sliced with a thin cuboid (thickness equal to slice thickness) using object. The object is then sliced with a thin cuboid (thickness equal to slice thickness) using object. The object is then sliced with a thin cuboid (thickness equal to slice thickness) using functions. The resulting intersection body is then positioned and ray traced POVRAYs Boolean functions. The resulting intersection body is then positioned an functions. The resulting intersection body is then positioned an . The positions of the light sources and the camera are adjusted as shown in the Fig. 2. The positions of the light sources and the cameraadjusted so asFig.

The positions of the light sources and the camera are adjusted as shown in the Fig. 2. The positions of the light sources and the cameraadjusted so asFig. 2: Concept of ray tracing.to obtain the slices at the correct scales. An "inside vector" was defined to identify the inside to obtain the slices at the correct scales. An "inside vector" was defined to identify the inside to obtain the slices at the correct scales. An "inside vector" was defined to identify the inside from the outside. When the scene is rendered, the camera shoots a ray for every pixel in the outside. When the scene is rendered, the camera shoots a ray for every pixel in the outside. When the scene is rendered, the camera shoots a ray for every pixel in the image to determine if it is inside the s lice object or outside thereby rendering the slice bitmap image to determine if it is inside the slice object or outside thereby rendering image.

The images obtained were then compressed using CCITT Fax4 scheme and sent to the image. The images obtained were then compressed using CCITT Fax4 scheme an image. The images obtained were then compressed using CCITT Fax4 scheme an DMD chip for exposure.III) Reconstructing Topology of TriaReconstructing Topology of Triangular Mesh:One of the major weaknesses of the STL file format is that it just contains a random list just contains a random list One of the major weaknesses of the STL file format is that it no particular order and does not contain topological information. If the topology topological information. If the topology of facets in no particular order edge connectivity, adjacency, no. of shells in the model) can be constructed edge connectivity, adjacency, no. of shells in the model) can be information (e.g. edge connectivity, adjacency, no.

of shells in the model) can be constructed edge connectivity, adjacency, no. of shells in the model) can be information (e.g. edge connectivity, adjacency, no. of shells in the model) can be be used to populate the ACIS data structure from bottom up can be used to populate the ACIS data structure from bottom up using information in the file, it can and thus could be used to translate the STL file into a format that ACIS can deal with. ACIS's d to translate the STL file into a format that ACIS can deal with. ACIS's d to translate the STL file into a format that ACIS can deal with. ACIS's STL file. Also, constructing this robust feature set can then be used to manipulate and slice the STL file.

ACIS's d to translate the STL file into a format that ACIS can deal with. ACIS's STL file. Also, constructing this robust feature set can then be used to manipulate and slice the STL file. Also, robust feature set can then be used to manipulate and slice information from the STL file will enable us to implement error correction information from the STL file will enable us to implement err topological data using information from the STL file will enable us to implement err and file manipulation routines which could prove to be very powerful in many applications. This and file manipulation routines which could prove to be very powerful in many applications. This and file manipulation routines which could prove to be very powerful in many applications. This was the premise in investing time in researchingthis method for slicing STL files.

This and file manipulation routines which could prove to be very powerful in many applications. This was the premise in investing time in researchingthis method for slicing STL files. time in researching this method for slicing STL files.able in literature to rebuild the There are a quiet a few data structures and algorithms available in literature to rebuild the There are a quiet a few data structures and algorithms avail topology information from a "bucket" of facets. Stephen Rock and Michael Wozny presented an topology information from a "bucket" of facets. Stephen Rock and Michael Woz topology information from a "bucket" of facets.

Stephen Rock and Michael Wozny presented an topology information from a "bucket" of facets. Stephen Rock and Michael Woz topology information from a "bucket" of facets. Stephen Rock and Michael Woz algorithm for extracting the adjacency information and for implementing basic error extracting the adjacency information and for implementing basic error extracting the adjacency information and for implementing basic error the computer graphics and visualization domain that There are others in the computer graphics and visualization domain that detection/correction [5]. There are others in address this issue [6] [7] [8]. Many of these algorithms use complex multidimensional data ]. Many of these algorithms use complex multidimensional data ]. Many of these algorithms use complex multidimensional data J Rossignac et al. structures for sorting three dimensional data which are difficult to implement. J Rossignac et al.

Many of these algorithms use complex multidimensional data J Rossignac et al. structures for sorting three dimensional data which are difficult to implement. J Rossignac et al. structures for sorting three dimensional data which are difficult to implement.of Georgia Tech proposed a simple integer array based data structure called the corner table data structure that is an elegant solution for working with triangular meshes [9, 10]. This data structure has been leveraged for our purpose and brief implementation details are discussed below.Corner Table:The Corner table data structure stores all of the topology and connectivity information in two simple integer arrays. The schematic in Fig. 3 shows the nomenclature and the integer arrays that hold the information. The region around a vertex in a facet is loosely referred to as a 'corner'. The vertex that corresponds to that corner is referred to as 'v(c)'. The corner opposite to the current corner 'c' is referred to as 'o(c)'.

The vertex that corresponds to that corner is referred to as 'v(c)'. The corner opposite to the current corner 'c' is referred to as 'o(c)'. The left and the right corners are respectively referred to as 'l(c)' and 'r(c)'. The next and previous corners are given by 'n(c)' and 'p(c)' respectively (assuming the vertices are listed in a counterclockwise manner). The triangle to which the current corner belongs is referred to as 't(c)'. [Refer to figure 3(a)]. The two integer arrays that store the connectivity information are the Vertex array 'V[c]' and the Opposite array 'O[c]' as shown in figure 3(c).For any given corner 'c', the corresponding vertex and opposite corner indices can be obtained from the Vertex array 'V[c]' and Opposite array 'O[c]' respectively. Once these two arrays are populated the adjacency information is available.

Once these two arrays are populated the adjacency information is available. For example, starting from a random corner 'c', we can access the left triangle by querying t(o(p(c))), the right triangle by t(o(n(c))) and the opposite triangle by t(o(c)). An edge array 'E[c]' can also be constructed in a similar manner to store the edge connectivity information. It would store the edge index of the edge opposite to a given corner 'c'.'V[c]' can be constructed by implementing a simple hash table and 'O[c]' can be constructed by identifying all the corners associated with a vertex and revolving around each vertex, marking the opposite corners. For more specific implementation details, refer to the original publication on corner table [9, 10]. Once the required arrays are populated, a simple command called 'swirl' can be easily implemented to identify the number of shells in the model.

Once the required arrays are populated, a simple command called 'swirl' can be easily implemented to identify the number of shells in the model. Refer to [9, 10] for more details.In this way, once all the required topological information is extracted from the STL file, it can easily be translated into a format that ACIS can recognize and work with. An algorithm with these ideas has been implemented and STL files were successfully sliced with the ACIS kernel.IV) Direct Slicing of STL files:An algorithm for directly reading and slicing STL files has also been implemented. One of the major issues in efficiently slicing an STL file is being able to quickly identify those facets that lie in the intersection region from the rest of the facets in the file. Tata et al. proposed a facet grouping strategy for this purpose [11]. Luo et al. proposed yet another strategy for identifying these facets quickly [12].

Tata et al. proposed a facet grouping strategy for this purpose [11]. Luo et al. proposed yet another strategy for identifying these facets quickly [12]. The algorithm presented here is very similar to theirs with minor modifications for speed improvements and to suit the needs of LAMP process. The schematic inFig. 4 illustrates the procedure for identifying the intersecting facets. In order to identify the intersecting facets, first each facet's maximum and minimum z-coordinates are computed and stored in memory (slicing assumed to be along the z-direction). Then, for a given slicing plane, first all the facets whose minimum z-coordinate is lesser than the slice plane height are selected (fig. 4(b)). Out of these selected facets, only those whose maximum z-coordinate is greater the slice plane height are identified and retained while the rest are discarded.

4(b)). Out of these selected facets, only those whose maximum z-coordinate is greater the slice plane height are identified and retained while the rest are discarded. This way, only those facets that intersect with the given slicing plane are isolated from the rest of the facets in the file (fig.4(c)).Fig.3 : Corner table data Structure: (a) Nomenclature. (b) sample triangular mesh with indexed corners, vertices and triangles. (c) Vertex table "V[c]" and Opposites Table "O[c]". (Note: in a manifold triangular mesh, there will not be any empty cells in the O[c] array as shown here. This is the case here because the array has been populated for a small portion of the mesh only).In order to do this, a data structure consisting of linked lists is implemented. It consists of a primary linked list sorted in the increasing order of z-values.

It consists of a primary linked list sorted in the increasing order of z-values. Each node in this linked list consists of its specific z-value and a pointer to a secondary list that contains all facets with the same minimum z-coordinate value as the z-value of that node. Once all the facets in the given STL file are populated in this data structure, it is straightforward to implement the rest of the operations required to accomplish the steps depicted in figure 4.(a)(b)(c)Fig. 4: Direct slicing of STL files: (a) Sample triangular mesh and slicing plane. (b) Facets with minimum z-coordinate lesser than the slice plane height selected. (c) Of the facets selected in (b), only those facets with maximum z-coordinate greater than the slice plane height are kept and the rest discarded.Once the facets in the intersection region are identified, a simple parametric intersection is computed between the facets and the slice plane to yield the various edges of the intersection wire.

Since the facets are listed in a random order, the wire edges are also computed in a random order. In conventional contour planning operations, these edges need to be sorted and the intersection loops need to be constructed. But for LAMP process, it is sufficient to generate a bitmap image of the slice. This can be directly accomplished by shooting rays for each row of pixels and computing intersection points and then using these points, the pixel values could be filled. This process of rasterizing the intersection wire is exactly the same as the one discussed in section I. The images thus obtained are saved in CCITT fax4 format and sent to the DMD chip for exposure.ResultsVarious data processing strategies for the LAMP technique have been investigated. A direct slicing algorithm has been implemented using ACIS kernel. This is the most preferred way of generating slice data for LAMP process because it doesn’t have the inherent inaccuracies associated with an STL file.

This is the most preferred way of generating slice data for LAMP process because it doesn’t have the inherent inaccuracies associated with an STL file. It is an efficient and accurate method to generate slice data and is especially suitable for LAMP process owing to the high feature resolutions. But through this approach, we can only handle specific CAD file formats ("SAT" file format or an equivalent format that is translatable to SAT). Since STL files are the defacto industry standard, several methods to process STL files have also been investigated. Slicing using POVRAY is a simple and elegant way to produce the data as it directly renders the bitmaps needed for LAMP process thereby alleviating all the intermediate steps that could lead to inaccuracies. But this approach is inherently slow owing to the underlying computationally intensive ray-tracing process. An alternate method to use ACIS for slicing STL files by reconstructing the topology information was investigated.

An alternate method to use ACIS for slicing STL files by reconstructing the topology information was investigated. An efficient data structure for extracting the topology information was implemented and the STL files were successfully translated to SAT files which were then processed using ACIS. This holds immense potential for implementing various STL file manipulation and error corrections routines using the robust inbuilt functionalities of ACIS. But the inherent drawback of this approach is that the resulting SAT files can be extremely large (several gigabytes) for complex STL files, because of the millions of entities in terms of vertices, edges and facets in the resulting CAD model and such enormous SAT files can be difficult to handle. Finally, to overcome these limitations, an efficient algorithm for directly reading and slicing STL files has been implemented.

Finally, to overcome these limitations, an efficient algorithm for directly reading and slicing STL files has been implemented. In conclusion, the ACIS approach was best suitable for direct slicing of CAD parts while the algorithm for directly reading and slicing triangular meshes yielded best results for STL files.References1. Guduri S, Crawford RH, Beaman JJ, "A method to generate exact contour files for solid freeform fabrication", Proceedings of the Solid Freeform Fabrication Symposium, Austin, Texas, pp 95–101, August 1992.2. Vuyyuru P, Kirschman CF, Fadel GM, Bagchi A, Jara-Almonte, " A NURBS-based approach for rapid product realization", Proceedings of the 5th International Conference of Rapid Prototyping, Dayton, Ohio, pp 229–239, June 1994.3. X. Chen, C. Wang, X. Ye, Y. Xiao and S. Huang, "Direct slicing from PowerSHAPE models for rapid prototyping", Int. J. Advanced Manufacturing Technology, 17(7), pp. 543–547, 2001.4. W. Cao, Y.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDIGITAL LIGHT PROCESSING (DLP): ANISOTROPIC TENSILECONSIDERATIONSE. Aznarte, C. Ayranci, and A.J. QureshiDepartment of Mechanical Engineering, University of Alberta, 10-203 DICE, 9211-116Street NW, Edmonton, AB, CanadaAbstractDigital light processing (DLP) 3D printing is an additive manufacturing (AM) process used to produce layered parts via photopolymerization. Anisotropy is a common characteristic of parts produced by DLP. Furthermore, printing conditions affect widely the resulting mechanical properties. This paper shows the effect of three printing factors on the final mechanical properties of specimens manufactured using DLP 3D printing. A series of ISO compliant tensile test specimens were designed, printed and tested.

A series of ISO compliant tensile test specimens were designed, printed and tested. The properties analyzed were the elastic modulus, ultimate tensile strength, ultimate strain and printing time. Preliminary findings on design guidelines for Vat Photopolymerization processes are presented in addition to the economic effect of the studied parameters in terms of the total printing time.Keywords: 3D Printing, Additive Manufacturing, Mechanical Characterization, Digital Light Processing (DLP), Vat-Photopolymerization, Stereolithography.1. IntroductionAdditive Manufacturing (AM) is the manufacturing process by which Three Dimensional (3D) parts are produced using an additive approach. Rapid prototyping was the first technology that allowed layer by layer material deposition using Computer Aided Design (CAD) (Wong and Hernandez, 2012). Stereolithography was the most popular rapid prototyping technique during the decade of 1990.

Stereolithography was the most popular rapid prototyping technique during the decade of 1990. It was invented in 1982 and patented in 1986 (Hull, 1986) by Charles Hull. This system builds a solid piece from a CAD model by sketching the cross section of the product with a laser beam that cures liquid photopolymer simultaneously (Comerford, 1993).Vat Photopolymerization is a form of AM. It uses a liquid bath of a polymeric photosensitive resin that is cured layer by layer through precise control and projection of a light source. The general composition of the resin is formed of a monomer material, a photoinitiator and an absorber (Gong et al., 2015). Photopolymerization is the process of linking monomers into polymers initiated by radiation exposure. This process can be divided in three steps: initiation, propagation and termination (Fuh et al., 1999).

This process can be divided in three steps: initiation, propagation and termination (Fuh et al., 1999). Said radiation, due to the exothermic nature of the process, can be a low-power UV laser or light source (Fuh et al., 1999).Parts manufactured by Vat Photopolymerization have varying mechanical properties depending upon the choice of parameters during the manufacturing. These properties depend mostly on the curing of the resin, which is affected by the power of the light source, the wavelength, the speed of construction, the resin composition and other factors (Salmoria et al., 2005). For example, a specimen subjected to a low exposure time and radiant power will be413under-cured; consequently, it would potentially have sub-par mechanical properties compared to a fully-cured specimen.Additive manufacturing differs from classical manufacturing processes in several aspects. Due to the layer by layer technique it ends up creating a material orthotropic in nature.

Due to the layer by layer technique it ends up creating a material orthotropic in nature. The AM process is usually fully digital, meaning that manufacturing process characteristics should be considered in the design stage, making the expertise of the designer a factor. The effect of the design decisions, such as part orientation or location on the base can affect the quality and properties of the part. Conventional design strategies are not sufficient to control the final properties of these parts. Consequently, with this study, we aim to achieve an early stage design method to understand and optimize the effects of printing parameters on elastic and mechanical properties of printed parts.In this work, we are reporting our findings on the effect of three chosen parameters on the final properties of printed specimens subjected to a tensile test. The variations between specimens will be analyzed and related to one or another parameter or the combination of them.

The variations between specimens will be analyzed and related to one or another parameter or the combination of them. Finally, the relationship between time and mechanical properties will be discussed, including the direct relation between time and cost. This work and the ongoing studies are crucial in establishing design curves for AM based materials represented in material selection charts.2. State of the ArtCertain combinations of printing parameters have been demonstrated to significantly improve elastic and mechanical properties of printed parts (Fatimatuzahraa et al., 2011). Unlike the bulk produced or processed conventional engineering materials, such as metals, or some polymers, the parts manufactured using AM processes demonstrate microstructural anisotropy.

Unlike the bulk produced or processed conventional engineering materials, such as metals, or some polymers, the parts manufactured using AM processes demonstrate microstructural anisotropy. The said behavior mostly originates from the layer by layer and sometimes directional nature of the production of the structure (Carroll et al., 2015).If the parts produced with AM are used without complete and fundamental understanding of this anisotropy, inconsistencies in properties can be observed (Santos et al., 2013). This, in return, will pose further challenges in design and simulation of AM parts with finite element analysis methods. Finally, another factor that may interfere with the mechanical properties is the scale at which each part is printed.

Finally, another factor that may interfere with the mechanical properties is the scale at which each part is printed. When testing similar parts produced by FDM processes with infills below 100% at different scales, the scalability of mechanical properties such as Elastic modulus (E), or Ultimate Tensile Strength (UTS), difference is not proportional to the scale of the piece(Mahmood et al., 2017; Qureshi et al., 2015).Extensive research literature is available on FDM process mechanical properties; however, there is very limited knowledge on such properties of photosensitive resins used in DLP 3D printing and the effect of the printing factors on them. Some researchers, like Stampfl et al. (2008), Liska et al (2007) and Schuster et al. (2007) focus on the chemical composition of the resin to modify the mechanical properties instead of focusing in the printing process itself.3.

(2008), Liska et al (2007) and Schuster et al. (2007) focus on the chemical composition of the resin to modify the mechanical properties instead of focusing in the printing process itself.3. Materials and Methodology414The investigation of design-material-process interaction for any process needs a thorough consideration of design aspects of the product being designed, the product raw material, and the manufacturing process needed to produce the part. Given the scope of this paper, the tensile response of a designed part is considered as the main design consideration. For this purpose, a standard design of a tensile specimen is adopted as per ISO 527 (Figure 1b). The primary processing technology is Vat Photopolymerization and the material used is the appropriate photopolymer resin.

The primary processing technology is Vat Photopolymerization and the material used is the appropriate photopolymer resin. The following sub-sections provide detailed information on the tools and materials used to carry out this study, as well as the methodology followed and the motivation to start the project.3.1 Materials and equipmentPrinter: the 3D printer used for this study is Ember DLP® by Autodesk. It uses a new 3D printing method in Vat Photopolymerization known as Digital Light Processing (DLP) (Gong et al., 2015). The main difference with classical stereolithography (SL) is that a light projector or Digital Micromirror Device (DMD) is used to define an individual layer at once by the polymerization of the photosensitive resin. A DMD is a semiconductor-based array of light switches. A light source is controlled by the switches using a binary code (Hornbeck, 1997).

A DMD is a semiconductor-based array of light switches. A light source is controlled by the switches using a binary code (Hornbeck, 1997). This results in a simultaneous, dual axis, 2D light projection on the cross-linking or polymerization plane as opposed to a sequential single laser beam scan normally used in SL systems.Ember has a resolution of 50µm in the plane parallel to the printing surface (XY resolution) and 10µm in the perpendicular plane to the printing surface (Z resolution). The Z resolution is determined by the layer height during the printing process. The printer's build maximum dimensions were reported as 64x40x134mm. The DLP is driven by glass optics and 0.45 WXGA DMD with a 912 x 1140 resolution array manufactured by Texas Instruments.Software: A solid modelling software was used to design the part; Autodesk Print Studio v1.6.5 was used as a primary computer aided manufacturing software (CAM) for slicing and creating the print files.

The primary control firmware for Ember was version 3.0. All data processing was performed using MATLAB R2016a.Resin Material: PR-48 type of resin was chosen for this study. This resin is also knownas Autodesk Standard Clear resin.Universal Testing Machine: The tensile tests were performed using Bose ElectroForce3200 Test Instrument.3.2 Methodology3.2.1Pre-processingEmber's software, Print Studio, provides a total of 64 parameters that can be modified. These parameters are divided into five main groups, namely: support, general, first layer, burn-in layer and model layer.

These parameters are divided into five main groups, namely: support, general, first layer, burn-in layer and model layer. "Support" refers to the extra material that must be added to make sure that the adhesion forces to the building surface are high enough to hold the part in place and to make415sure that every section of the part is connected to said surface at every moment during the printing process.– Under the "general" section there are three settings to check, namely; anti-alias, grey scaleand activation or deactivation of the homing sensor.– "First layer" parameters are those that refer to the printing settings of the first layer.– "Burn-in layer" refers to the settings of a number of layers right after the first one, typically less than 10.

These layers usually are printed at lower speeds and longer exposure time than the model layers to reaffirm the contact with the building platform.– "Model layers" are all the layers left after the last burn-in layer.Due to different layer types during the part fabrication, 16 of the parameters are repeated within the 3 layer groups. On top of the settings offered by the software, other factors such as location, orientation and rotation of the printed parts can be identified, making it a total of 67 factors. From this point of the study, every printer parameter or software setting will be designated as a "factor".The experiments carried out for this study are a combination of 3 controllable factors out of the 67 identified. These factors were chosen based on the most common controllable factors in DLP printing as a preliminary study to prove the variation in mechanical properties when printing factors are changed.

These factors were chosen based on the most common controllable factors in DLP printing as a preliminary study to prove the variation in mechanical properties when printing factors are changed. Each factor was assigned 3 levels or values that are summarized in Table 1.– Factor number 1, layer thickness, represents the total displacement of the Z-axis betweeneach light exposure.– Factor number 2, exposure time, is the total time that the blue light (405 nm wavelength)is turned on projecting the desired shape over the uncured resin.– Factor number 3, part orientation presents the alignment of the specimen with the printer'saxis.The physical representation of each orientation level can be observed in Figure 1.Table 1. Factors and levelsFactor Layer thickness (µm) Exposure time Part orientationLevel 1 10 1.6 XLevel 2 25 1.8 YLevel 3 50 2.0 ZThe material response was assessed using mechanical tests.

Factors and levelsFactor Layer thickness (µm) Exposure time Part orientationLevel 1 10 1.6 XLevel 2 25 1.8 YLevel 3 50 2.0 ZThe material response was assessed using mechanical tests. The tensile test specimens used were built and tested using Type 1BB specimens of ISO-527-1/2:2012 (ISO527 standard). The specimen type was selected due to the limitation of the maximum load that the testing apparatus can handle during a test. A schematic view of the print patterns and the technical drawing of the specimens can be found in Figure 1. The model was created in SOLIDWORKS®4162015 and then saved as an STL file. The same part was imported 5 times to Print Studio; all the copies were distributed on the printing surface following the same orientation.Figure 1.

The same part was imported 5 times to Print Studio; all the copies were distributed on the printing surface following the same orientation.Figure 1. (a) Part orientation levels on the printing surface, and (b) Design of ISO-527-1/2:2012 Type 1BB Test Specimen (ISO, 2012)Five specimens were prepared for each of the settings mentioned above; resulting in a total of thirty specimens distributed in 6 combinations of levels. Depending on the parameters chosen for each batch, a resulting total print time was obtained. Each print was determined to be representative of one or two print parameters; indicated in bold font in Table 2.All the specimens were printed in groups of five in order to establish the repeatability of the process; this number was chosen because 5 is the maximum number of specimens that could be fit at the same time on the building surface aligned with the X axis, which was taken as the limiting orientation.

It must be noted that the total print time is independent of the number of specimens printed on each batch as long as the orientation is kept constant.Table 2. Printing factorsPrint Number Print Time Layer Thickness (μm) Exposure Time (s) Part Orientation1 1:29 10 1.6 Y2 0:35 25 1.6 X3 0:19 50 2 Y4 1:51 50 1.6 Z5 0:33 25 1.8 X3.2.2Post-processingDirectly after the completion of printing process, each part was individually detached from the build surface and immersed in isopropyl alcohol for 3 minutes as recommended by the manufacturer of the machine. Alcohol dissolves any uncured resin and cleans the surface of the parts. Following this, the specimens were left to dry on a clean surface for 24 hours before tested.The tensile test was performed at a constant speed of 0.01mm/s. The grips were placed at an initial distance of 25 mm from each other. The load cell used for testing the samples was rated4176 0:17 50 2 Xfor a maximum force of 400 N.

The grips were placed at an initial distance of 25 mm from each other. The load cell used for testing the samples was rated4176 0:17 50 2 Xfor a maximum force of 400 N. The properties that are investigated were calculated according to the ISO-527:2012 standard (ISO, 2012).4. MotivationBased on the previous studies on FDM (Rahman et al., 2015) we can state that a specimen built with layers perpendicular to the direction of the force, as observed in Figure 2, will fail at a lower stress than a specimen printed with layers parallel to the force direction.Figure 2. Layers result of printing in Z directionWith this study, we aim to find out the best strategies of part placement on the build surface to maximize the ultimate strength. As well, we would like to establish a method to predict the mechanical properties of any part printed with the resin PR-48 with any print setting.

As well, we would like to establish a method to predict the mechanical properties of any part printed with the resin PR-48 with any print setting. To do so, we will analyze the tensile test results of the tested parts, focusing on elastic modulus, ultimate tensile strength and ultimate strain. We will highlight the most significant findings and propose possible justifications.The total print time will also be analyzed and compared to the tensile test responses.Being time directly related to cost, every finding associated to time can be extended to cost.Thirty samples were tested at 6 unique combinations of factors. The average elastic modulus, ultimate stress and ultimate strain of the five repetitions for every test are represented in the following section. Each result is grouped and compared among the controllable factors.5. Results and discussionIn this section, the results of the tensile test will be shown and discussed.

Each result is grouped and compared among the controllable factors.5. Results and discussionIn this section, the results of the tensile test will be shown and discussed. The section is divided into three subsections discussing the effect of different factors on: elastic modulus, ultimate stress, ultimate strain and time respectively.5.1 Tensile Test Results - Elastic modulusThe average elastic modulus of all the specimens tested was 0.749 GPa with a standard deviation of ±0.196 GPa. However, by differentiating the results in terms of the printing factors it can be seen that the moduli follow a trend. This can be observed in Figure 3 where the moduli results are grouped under the four major categories, layer thickness, exposure time, part orientation and printing time.Elastic modulus vs.

This can be observed in Figure 3 where the moduli results are grouped under the four major categories, layer thickness, exposure time, part orientation and printing time.Elastic modulus vs. layer thickness: specimens printed with a layer thickness of 10μm, 25 μm, and 50 μm have an average elastic modulus of 1.031 GPa, 0.719 GPa and 0.651 GPa with a standard deviation of ±0.028 GPa, ±0.046 GPa and ±0.020 GPa, respectively. As can be seen,418increasing layer thickness from 10 µm to 50 µm has decreased the modulus a 37% (from 1.031 to 0.651 GPa). The standard deviation for 25 µm is greater than the others, making the result significantly closer to the 50 μm's than to the 10 μm's. Results also suggest that the smaller the layer, the higher the elastic modulus.

Exposure time, (c)Elastic modulus vs. Part Orientation, (d) Elastic modulus vs. Printing Time.Elastic modulus vs. exposure time: the inverse trend is observed for the exposure time. Less exposure time implies a lower resulting elastic modulus. This tendency could be associated to the curing of the resin; low curing times implies a less stiff material, while a longer curing time results in a stiffer specimen. In this case, the difference in results is higher between 1.6s (with an average elastic modulus of 0.426 GPa) and 1.8s (with an average elastic modulus of4190.772 GPa) than between 1.8s and 2.0s (with an average elastic modulus of 0.898 GPa). Following the reasons exposed for the layer thickness, a possible explanation for the high elastic modulus for longer exposure times is that more light gets to the previous layer, assuring the adhesion between them.Elastic modulus vs.

part orientation: the lowest elastic modulus is observed for the specimen printed in the Z direction, being in average, 0.426 GPa with a low standard deviation of ±0.018 GPa. On the other hand, higher elastic moduli are obtained from the X and Y directions, with an elastic modulus of 0.719 GPa and 1.031 GPa respectively. This result confirms that the lower mechanical properties are those of parts printed along the Z direction. However, X and Y directions should show alike properties due to the printer’s configuration.Elastic modulus vs. printing time: the printing time varies between 17 minutes and 1 hour and 51 minutes, as can be observed in Figure 3(d). In this range of time, the average elastic modulus goes from 1031 MPa to 426 MPa. It should be noted that a 25% increase of printing time (1 hour and 29 minutes compared to 1 hour and 51 minutes) leads to a decrease of almost 60% in elastic modulus (1.031 GPa with respect to 0.426 GPa).

Finally, as observed in Figure 3(d), where the bars are ordered from short print times to long print times from left to right, long printing times are not directly related to higher elastic modulus.5.2 Tensile Test Results - Ultimate stressThe average ultimate stress of the 30 tests performed is 21.7 MPa, with a standard deviation of ± 6.5 MPa. The summarized average ultimate stresses can be observed in Figure 4. In a similar way to the elastic modulus, the ultimate stress follows some trends when sorting the results per the factors used.Ultimate stress vs. layer thickness: specimens printed with the thinnest layer thickness can withstand greater forces than those printed with thicker layers, see Figure 4(a). This result can be associated to the resin transmittance, which provides higher degree of curing to a thin layer than to a thicker layer.

This result can be associated to the resin transmittance, which provides higher degree of curing to a thin layer than to a thicker layer. However, in this case the differences between the 10μm and 25μm can be almost neglected considering the standard deviation of each average ultimate stress, being 26.3 ± 1.4 MPa and 23.0 ± 1.3 MPa. The same effect occurs between 25μm and 50μm, being the average and standard deviation 19.9 ± 0.9 MPa.Ultimate stress vs. exposure time: the results when comparing the exposure times, seen in Figure 4(b), are much more differentiated than the previous ones. Parts whose layers have been cured during 1.6 seconds appear not to be fully cured; hence the average ultimate stress is about three times lower (8.7 MPa) than the stresses under which specimens exposed for 2 seconds break (27.7 MPa).Ultimate stress vs. part orientation: the part orientation plays an important role in the ultimate stress result.

part orientation: the part orientation plays an important role in the ultimate stress result. Specimens printed along the Z direction show a low ultimate stress. This fact can be related to the adhesion forces between layers oriented perpendicularly to the tensile force direction.420Ultimate stress vs. printing time: looking at Figure 4(d) one result must be highlighted; the longest printing time results in an extremely low ultimate stress, only 8.7 MPa; while all the other printing times offer average ultimate stresses between 19.8 MPa and 27.7 MPa.

Printing Time.5.3 Tensile Test Results - Ultimate strainThe average ultimate strain of all specimens tested was 5.8% a with a standard deviation of ±1.87%. In Figure 5, ultimate strain results are plotted and grouped in terms of the factors analyzed and the total printing time. In general, the results obtained when looking at the ultimate421strain of the specimens, Figure 5, are drastically different than the ones shown previously.

(a) Ultimate strain vs. Layer Thickness, (b) Ultimate strain vs. Exposure Time, (c)Ultimate strain vs. Part Orientation, (d) Ultimate strain vs. Printing Time.Ultimate strain vs. layer thickness: by observing Figure 5(a) the lowest output is obtained with specimens printed with 10µm layers, being the ultimate strain 4.8%; layers of 25µm and 50µm offer ultimate strain results of 7.0% and 5.6% respectively. These last results, when accounting for the standard deviation, overlap with each other; hence, ultimate strains of 25µm and 50µm layers is not so relevant compared to other factors.Ultimate strain vs. exposure time: the exposure time, found in Figure 5(b), brings almost identical ultimate strains for 1.8s and 2.0s (7.33% and 7.27% respectively) suggesting a less422brittle structure than the specimens with only 1.6s of exposure time (ultimate strain of 2.72%).

This high difference is due to the higher curing degree of the layers exposed during a longer time; while at 1.6 seconds, the curing degree is lower.Ultimate strain vs. part orientation: parts printed in the Z direction failed at lower strains (2.72%) compared to those printed in the other two directions (7.02% for X and 4.79% for Y). The significant difference between X and Y direction, seen in Figure 5(c) is due to the other factors applied when printing the parts or noise coming from humidity or other sources; given that, theoretically, there should be low to no difference between them.Ultimate strain vs. printing time: there is not a clear relationship between total printing time and ultimate strain, shown in Figure 5(d). The lowest ultimate strain, 2.72%, is the result of the highest printing time, 1hour and 51 minutes. While the one of the highest ones, 7.27%, corresponds to the lowest printing time, 17 minutes.

The lowest ultimate strain, 2.72%, is the result of the highest printing time, 1hour and 51 minutes. While the one of the highest ones, 7.27%, corresponds to the lowest printing time, 17 minutes. Again, the main reason is that the longest print is for specimens oriented along the Z axis and at low exposure times.6. ConclusionsManufacturing processes are an important link of the design-material-process chain and impact the final quality and properties of the part significantly. The considerations or effect of manufacturing process on the product is undertaken during the Design for X (DfX) activities in the product design. In case of traditional/subtractive manufacturing processes, a vast body of design guidelines are available and integrated by product designers in design phase. However, as shown in this paper, there is very limited understanding of the design strategies for Additive Manufacturing processes, specifically for Vat Photopolymerization process.

However, as shown in this paper, there is very limited understanding of the design strategies for Additive Manufacturing processes, specifically for Vat Photopolymerization process. This is complicated by the fact that, unlike the traditional manufacturing processes, the AM processes also exhibit orthotropic behavior; which must be considered while carrying out product design. In order to address this mechanical behavior, the characteristics of the process must be taken into account when carrying out the parametric design of the product.In this paper, we report on the preliminary findings of ongoing efforts on developing design for additive manufacturing guidelines for Vat Photopolymerization process with a focus on print time, cost, elastic modulus, ultimate tensile stress and ultimate strain. As this paper shows, significant variations between mechanical properties exist for the same designed part printed with different levels of manufacturing process factors.

As this paper shows, significant variations between mechanical properties exist for the same designed part printed with different levels of manufacturing process factors. It was shown that variations in the levels of layer thickness, exposure time and part orientation resulted into significant variations in the part quality. A 59% difference in elastic modulus was observed when parameters were changed; this variation was calculated taking the average results. The difference in ultimate stress using different exposure times was 68%. Finally, the maximum strain was proven to change by 63% when the parameters are changed. It was also shown that merely changing the orientation of the part in the printer has a significant effect on the strength of the part.In addition to the mechanical properties, this paper also sheds light on the economic aspects of the Vat Photopolymerization in terms of time taken to print and its relationship with the mechanical properties.

The time needed to print is an important factor representing the machine operating time, hence linked directly to cost of running, depreciation, as well as423maintenance. The paper shows that by varying the factors, the time to print the test sample varies between 17 and 111 minutes, representing an increase of 6.5 times. Moreover, in case of tensile strength the best strength is obtained at the lowest build time.Considering the results from this study, a more detailed study is being carried out to design and manufacturing parameter optimization for the mechanical properties. This includes an extended design of experiments study to establish the correlation and interaction between different factors.

This includes an extended design of experiments study to establish the correlation and interaction between different factors. This will result into a comprehensive framework for mechanical properties in function of the printing factors, allowing printing of functional parts; making the DLP printing a viable additive manufacturing process when building specimens that are going to be subjected to external loads.ReferencesCarroll, B.E., Palmer, T.A. and Beese, A.M. (2015), “Anisotropic tensile behavior of Ti-6Al-4V components fabricated with directed energy deposition additive manufacturing”, Acta Materialia, Acta Materialia Inc., Vol. 87, pp. 309–320.Comerford, R. (1993), “A quick look at rapid prototyping”, IEEE Spectrum, Vol. 30 No. 9, pp.28–29.Fatimatuzahraa, A.W., Farahaina, B. and Yusoff, W.A..

18 No. 12, p. 125014.Wong, K. V and Hernandez, A. (2012), “A Review of Additive Manufacturing”, ISRNMechanical Engineering, Vol. 2012, pp. 1–10.AcknowledgmentsThe authors would like to acknowledge the funds provided by Natural Sciences and Engineering Research Council of Canada (NSERC) - Discovery Grants (418533 and RGPIN- 2016-04689) and the University of Alberta Mechanical Engineering Department that allowed this research to be conducted. The authors also would like to acknowledge the funds provided by Canadian Foundation for Innovation (CFI - Project #31500).Permission to use extracts from ISO 527-2:2012 was provided by the Standards Council of Canada (SCC). No further reproduction is permitted without prior written approval from SCC.425.

GITAL MICROFLUIDIC DESIGN FOR INJECTION CONTINUOUS LIQUID INTERFACE PRODUCTION OF 3D OBJECTSGabriel Lipkowitz1, Tim Samuelsen1, Kaiwen Hsiao2, Maria T. Dulay2, Ian Coates3, William Pan1, Eric S.G. Shaqfeh1,3, Joseph M. DeSimone2,31 Department of Mechanical Engineering, Stanford University, Stanford, CA 94305 2 Department of Radiology, Stanford University, Stanford, CA 94305 3 Department of Chemical Engineering, Stanford University, Stanford, CA 94305AbstractIn additive manufacturing, it is imperative to increase print speeds, use higher viscosity resins, and print with multiple different resins simultaneously. To this end, we introduce a new UV-based photopolymerization 3D printing process exploiting a continuous liquid interface—the deadzone—mechanically fed with resin at elevated pressures through microfluidic channels dynamically created and integral to the growing part.

Through such mass transport control, injection continuous liquid interface production, or iCLIP, accelerates printing speeds 5 to 10-fold over current methods such as continuous liquid interface production (CLIP), can utilize resins an order of magnitude more viscous than can CLIP, and can readily pattern a single heterogeneous object with different resins in all Cartesian coordinates.

We characterize the process parameters governing iCLIP and demonstrate use-cases for rapidly printing carbon nanotube-filled composites, multi-material features with length scales spanning several orders of magnitude, and lattices with tuneable moduli and energy absorption.Keywords: injection continuous liquid interface production (iCLIP), continuous liquid interface production (CLIP), generative design, parametric design, algorithmic design, additive manufacturing, microfluidicsIntroductionUV curable liquid resin-based additive manufacturing (AM) can broadly be divided into three generations of vat photopolymerization (1) and material jetting. In first generation vat photopolymerization (VP), also known as stereolithography (SLA), a bath containing a single liquid resin is photocured spatio-selectively by a scanning laser point source.

In first generation vat photopolymerization (VP), also known as stereolithography (SLA), a bath containing a single liquid resin is photocured spatio-selectively by a scanning laser point source. Second generation VP, referred to as digital light projection (DLP), uses a rapid sequence of projected UV images that span the entire X-Y plane of a bath containing a liquid resin in a single exposure. Often referred to as third generation VP, continuous liquid interface production (CLIP, Figure 1a) relies on resin renewal at the build surface through a continuous liquid interface—the dead zone—created by oxygen, a polymerization inhibitor, fed through the oxygen-permeable window beneath the vat (2-4). CLIP, achievable with multiple patterns of platform movement and UV exposure (5) along with different window configurations (6,7), enables printing at speeds up to 3000 mm/hr, 25 to 100 times higher than traditional AM methods.

While to date CLIP has been limited to relatively low viscosity resins (commercially available resins from Carbon, Inc. have viscosities up to roughly 2500 cP) (8), CLIP produces isotropic parts, unlike conventional 3D printing methods like fused filament fabrication (FFF) and powder bed fusion (PBF), and has proven suitable for manufacturing at high volumes and at high resolution for e.g. biomedical devices (9,10).1969However, CLIP is still significantly slower than injection molding. This is due to severe mass transport limitations on resin flow through the thin dead zone (Figure 1b), inducing Stefan adhesion forces that require delay time to equilibrate the negative dead-zone pressure. These forces also limit print sizes and necessitate cumbersome supporting scaffolds as in traditional VP techniques (11,12).

These forces also limit print sizes and necessitate cumbersome supporting scaffolds as in traditional VP techniques (11,12). From lubrication theory, this Stefan adhesion force scales dramatically with part radius:FStefan=3πR4μU 2h31)(where R is part radius, μ resin viscosity, U print speed, and h dead zone thickness. These mass transport limitations, though intensely examined in the literature (13-15), restrict CLIP and VP from efficiently processing highly filled composite resins, and from effectively printing with multi-materials.In the second UV resin-based AM approach, material jetting (MJ), an inkjet print head deposits photocurable resin droplets onto a build surface, readily printing with high spatial control, multi-materials, and highly viscous filled composites (16-18).

Disadvantages of MJ, however, include highly anisotropic part properties and very slow printing, making it difficult to scale and arguably mainly for prototyping.Here we describe our new AM method that synthesizes these two established AM approaches by injecting resin through CAD software-designed microfluidic conduits in the part, or “viaducts”, into a continuous liquid interface to supply part production (Figure 1b). Like CLIP, this method uses a highly oxygen permeable window to create a deadzone that, in this instance, provides a destination for resin flow from our high pressure fluidic injection system (Figure 1c). We call this 3D printing approach injection continuous liquid interface production, or iCLIP for short.Results and DiscussionInjection into a deadzone alleviates suction forces to accelerate printing of 3D geometriesAn apparatus for testing the iCLIP printing process is described in Figure 2a.

In this system, without high-pressure injection, we recapitulate a traditional CLIP process, and suction forces scale with part cross sectional area, as expected from the aforementioned lubrication theory (Figure 2b). Without limiting volumetric print speeds, significant defects and premature delamination from the platform can result.Herein we show that a single viaduct can offset these forces. From lubrication theory, one can determine that a single viaduct, whose area can be varied dynamically by software during printing, introduces a positive pressure increment by administering resin at a volumetric flow rate Q, according to:pviaductμLQ Ah3(2)∼where μ is resin viscosity, L part length, A viaduct area, and h dead zone thickness. Dead zone thicknesses in iCLIP are very small, as in traditional CLIP and as expected from analytical models (19), typically on the order of tens of microns.

(b) Injection CLIP process indicating the flow of injected resin from a pressurized source through microfluidic ducts engineered within the growing part into the dead zone. (c) Analytically-derived dead zone velocity fields from lubrication theory during traditional CLIP (left) and Injection CLIP with continuous injection at two different1971volumetric flow rates through a central viaduct (center and right), where z and r are vertical and radial distances in the deadzone, respectively, and vr is radial velocity.We find this positive pressure increment indeed allows iCLIP to significantly reduce Stefan adhesion forces—for large area parts by almost two orders of magnitude (Figure 2c). While we still observe some scaling of measured Stefan adhesion forces with part cross sectional area, this scaling is much less dramatic than in traditional CLIP.Figure 2: iCLIP alleviates suction forces that plague the traditional CLIP process.

(a) Design of a prototype iCLIP printer, with overall hardware configuration and salient features of the iCLIP system highlighted. In upper left, the build region of the iCLIP printer, with the mounting of the load cell and fluidic connectivity ports indicated. The downwards Stefan adhesion force measured by the load cell is indicated by the black arrow, while the platform movement is indicated in grey. (b-c) Experimental load cell force data measured for three consecutive layers, each of 3 second duration, while printing a conical geometry with varying cross sectional areas by CLIP and iCLIP, respectively.These resin injection ducts enable a significant reduction in suction forces that in turn allows iCLIP to achieve significant increases in print speed. For both CLIP and iCLIP, maximum printable speeds for a given cross sectional area can be quantified as the part draw rate at which delamination events occur, at statistically significant levels, for multiple primitive geometries.

Injecting resin through a single central viaduct, guided by computational fluid dynamic simulations (Figure 3a-d) we observe increases in the maximum achievable print rates determined to be between 5 and 10-fold (Figure 3e-h) over CLIP. For a solid conical part with a cross sectional radius growing from 2 mm to 20 mm, for instance, the maximum linear print speed with traditional CLIP before delamination occurs is ~20 mm/hr. With iCLIP, by1972contrast, maximum print speeds up to ~125 mm/hr are achievable. With a single resin injection channel, delamination does still occur at an elevated print speed for a given part cross sectional area due to the difficulties in administering sufficient resin flow through a single duct, a shortcoming that can be overcome by more complex duct geometries, which is a topic of ongoing investigation.Figure 3: Injection accelerates printing of solid 3D geometries by CLIP.

(a-d) Computational fluid dynamics predictions of the pressure fields experienced while printing a solid circular cross section by traditional CLIP without injection (a), by injection CLIP with flow through a central duct at varying volumetric flow rates (b-c), and by injection CLIP with flow through four distributed ducts at equal flow rates (d). (e-h) Quantified maximum print volumetric throughputs for two test geometries with varying cross sectional areas, cone (e, g) and cylinder (f, h), by CLIP (left) and iCLIP (right). Grey dotted lines indicate delamination- free prints. Error bars denote +/- one standard deviation from three independent print trials.Importantly, such ducts need not have detrimental impacts on the mechanical propertiesof the final part with appropriate post-processing. In this study, viaduct channels were1973always sealed after printing by utilizing a post-printing UV cure of the resin-filled channels.

In this study, viaduct channels were1973always sealed after printing by utilizing a post-printing UV cure of the resin-filled channels. When such standard protocols are followed, we do not observe any difference in mechanical properties between CLIP and iCLIP printed dogbones, as shown in Figure 4.The requirement to integrate ducts into a part can impact the resolution of iCLIP incomparison with CLIP traditionally, but with careful design strategies, this can be minimized. In areas of the part where ducts must be included, feature resolution does drop to the minimum achievable channel diameter before concerns arise such as capillary collapse, if the channel is freestanding, or channel cure-through, if the channel is embedded.

Moreover, viaduct channels can be designed to have larger radii than this achievable minimum, because of the challenges associated with enforcing viscous flow through a narrow channel, according to the Hagan-Poiseulle equation:(5)8𝜇𝜇𝐶𝐶𝑄𝑄 4Δ𝑃𝑃 =𝜋𝜋𝑅𝑅where P is pressure, μ dynamic viscosity, L channel length, Q flow rate, and R channel radius. Flow rates typically range, in this study’s experiments, from 7 μL/s to 27 μL/s.While such channels can therefore be detrimental to iCLIP resolution in some circumstances, it is important to note that ducts need not be engineered into all regions of an iCLIP part. Here, the native feature resolution of a traditional CLIP process generalizes to this new printing platform, along with the same tradeoff between print speed and feature resolution as characterized by its inventors (2).

Given that the smallest features are indeed those that experience the smallest Stefan adhesion forces, which thus do not likely require active injection to offset, such high resolution features can be preserved in iCLIP parts. Thanks to the software- guided nature of the integration of ducts into iCLIP parts, moreover, ducts can be designed to avoid such high resolution features.1974Figure 4: Negligible impact of post-cured viaduct on iCLIP printed part mechanical properties. (a) CAD design of test geometry for evaluating effect of viaduct on dogbone mechanical properties. (b) Picture of iCLIP printed dogbone, with inset showing optical microscopy image of central viaduct. (c) Comparison of iCLIP printed dogbone with ASTM Type V specifications. (d-e) Ultimate tensile strength and Young’s modulus, respectively, of iCLIP printed dogbones under different viaduct post-treatment conditions.

(d-e) Ultimate tensile strength and Young’s modulus, respectively, of iCLIP printed dogbones under different viaduct post-treatment conditions. All measurements represent the average of three independently printed dogbones, with error bars denoting +/- one standard deviation from the mean.ConclusionIn summary, we describe here iCLIP, a novel 3D printing method using active control of mass transport during continuous liquid interface printing to enhance print speeds. Ongoing work in optimizing the existing iCLIP process focuses on detailed modeling of the flow boundaries in the deadzone, optimizing flow rates to minimize Stefan adhesion forces and cavitation, and automating the generation of viaduct geometries and injection profiles to accelerate iCLIP printing.

Future work in extending iCLIP to new materials and geometries will focus on testing a wider library of viscous filled resins with superior mechanical and electrical properties for applications in smart and sensor-embedded product designs, along with developing predictive models for analyzing the mechanical properties of multi-material iCLIP structures for applications in 4D printing and soft robotics, among other areas.1975Materials and Methods:Design of a prototype iCLIP printerFor print platform motion, a NEMA 57 stepper motor supplied by a 12V power bankwas used to drive vertical build platform translation along a 12-inch Stroke Linear Motion Router (VXB Ballbearings, Anaheim, CA, USA). The UV light engine used was a 3DLP9000 (Digital Light Innovations, TX, USA) with a 4 million pixel 2560 ✕ 1600 digital micromirror device (DMD), configured with a 385 nm LED and a 30 μm field of view (FOV) projection lens, with a total projection area of 76.8 mm x 48 mm.

The light engine is a combination of a digital micromirror device (DMD) chip set (DLP9000, Texas Instrument, TX) along with a projection lens; the intrinsic specification of the DMD chipset is 385 nm ultraviolet (UV) wavelength, 2560 ✕ 1600 DMD array, 7.6 μm ✕ 7.6 μm pixel size, build area of 19.5 mm ✕ 12.2 mm; the projection lens diverges the UV projection to a 2560 ✕ 1600 array of 30 μm ✕ 30 μm pixels to a build area of 76.8 mm ✕ 48.0 mm at a working distance of 126.5 mm. The printer was coordinated with an Arduino MEGA 2560 microcontroller and RAMPS 1.4 shield running open-source Marlin firmware.

The printer was coordinated with an Arduino MEGA 2560 microcontroller and RAMPS 1.4 shield running open-source Marlin firmware. Custom software, written in C++ and implemented in the Qt Integrated Development Environment to provide a graphical user interface, allowed for tailoring UV light intensity, UV exposure time, stage speed and acceleration, layer thickness, and delay time after layers, within and between prints.Resin formulationsFor print speed and resin viscosity experiments, resins of tuneable viscosity wereprepared by mixing isobornyl methacrylate, bisphenol A ethoxylated acrylate, and bisphenol A glycidyl methacrylate at varying ratios with 0.7 weight percent phenylbis(2,4,6- trimethylbenzoyl)phosphine oxide and 0.06 weight percent UV absorber 2-tert-butyl-6-(5- chloro-2H-benzotriazol-2-yl)-4-methylphenol (BLS 1326), all from Sigma Aldrich (St. Louis, MO, USA), using a Thinky planetary mixer (Thinky USA, Inc., Laguna Hills, CA, USA).

Louis, MO, USA), using a Thinky planetary mixer (Thinky USA, Inc., Laguna Hills, CA, USA). Elastomeric resin formulation was prepared with varying ratios of epoxy aliphatic acrylate (EAA, tradename Ebecryl 113) and aliphatic urethane-based diacrylate (AUD, tradename Ebecryl 8413), which were purchased from Allnex (Malaysia), diluted in isobornyl acrylate and mixed with 1.0 weight percent photoinitiator diphenyl(2,4,6-trimethylbenzoyl)phosphine oxide (TPO).Load cell measurementsThe build platform was designed to accommodate a Miniature S-Beam Jr. Load Cell 2.0 (Futek, Irvine, CA, USA) of dimensions 0.75" x 0.69" x 0.26", with a resolution of ±0.05 percent, a rated output of 1 mV/V (250 g) to 2 mV/V (1 – 100 lb), a bandwidth of 2,000 cycles per second, and with signal processing via a USB Load Cell Digital Amplifier (Futek, Irvine, CA, USA). Measurements were taken at 100 Hz.

Measurements were taken at 100 Hz. Force data was obtained midway through the print, i.e., once the build platform had fully exited the resin vat such that buoyancy forces did not change between layers.1976Computational Fluid Dynamics SimulationsTo control multi-material iCLIP printing, the flow within the deadzone was simulatedusing commercially available CFD software (ANSYS Fluent, Canonsburg, PA, USA). The resin was simulated as a homogenous fluid with a non-Newtonian viscosity profile. The fluid was given a density of 1120 kg/m3. The computational domain consisted of the vat resin and deadzone, modelled as fluid, and the printed object, as solid. The domain was discretized using an unstructured, hexahedral cell mesh comprised of a Cartesian core mesh. A no-slip boundary condition was prescribed on the deadzone-window interface. At the viaduct inlet of the computational domain, a mass influx profile is applied corresponding to the prescribed flow rate.

At the viaduct inlet of the computational domain, a mass influx profile is applied corresponding to the prescribed flow rate. The steady-state solution of the flow within the iCLIP printer is obtained by solving the conservation of mass and momentum equations on the computational mesh. Meshes were generated with an element size of 0.01 mm.

Meshes were generated with an element size of 0.01 mm. Simulations were either run as static, when validating against injection calibration experiments, or dynamic, when guiding multi-material prints.Mechanical TestingBefore testing, prints were post-processed by washing with 99% isopropanol, wickingaway excess resin with Kimwipes, and post UV curing either by irradiating in an APM LED UV CUBE II oven (APM Technica, Switzerland, 365 nm), or with a handheld Loctite CureJet UV LED Controller (Henkel Corp., Dusseldorf, Germany).Tensile testing of iCLIP-fabricated ASTM D638 type V dog bones, assessed fordimensional accuracy using a micrometer, was conducted using an Instron 5566 (Universal Testing Systems, Stoneham, MA, USA) with a cross-head speed of 1 mm/min at 25 °C to achieve the break at roughly 60 s, which is in accordance to the 30 s – 5 min outlined in ASTM D638, with a 100 N load cell.

Tensile strength was calculated using the maximum load of the stress/strain curve and Young’s modulus using the linear portion of the stress/strain curve.Mechanical properties tests of composite prints were performed with an MTS Criterion Model 42 Universal Testing machine (MTS Systems Corporation, Eden Praire, MN, USA), equipped with a 100 N load cell and either with fixtures for 3-point bending or platens for uniaxial and transverse compression. For uniaxial cyclic compression tests, the specific energy absorption of the cylindrical specimens (in J/kg) was calculated as the energy dissipated (from the area between the loading and unloading curves in the hysteresis loop) per unit mass.1977Acknowledgements: Imaging experiments for this work were performed at the Molecular Imaging Program at Stanford (MIPS). Mechanical testing was performed at the Blume Center for Earthquake Engineering at Stanford.

Mechanical testing was performed at the Blume Center for Earthquake Engineering at Stanford. For printer design and prototyping, part of this work was performed at the Biodesign Collaboratory, supported by the Stanford Byers Center for Biodesign. For rheological characterization, part of this work was performed at the Stanford Soft Matter Facility (SMF).Funding: This work is supported in part by the Precourt Institute for Energy at Stanford University. GL is supported by a graduate fellowship from the National Science Foundation.Author contributions: GL built the printer, performed experiments, executed simulations, illustrated schematics, and wrote the manuscript. TS wrote printer software. ES, GL, and KH formulated lubrication theory. GL designed the printer. GL, IC, MTD, and WP developed resin formulations. WP and GL took photographs. All edited the manuscript.

ES, GL, and KH formulated lubrication theory. GL designed the printer. GL, IC, MTD, and WP developed resin formulations. WP and GL took photographs. All edited the manuscript. JMD conceptualized and directed the project.Competing interests: JMD declares that he has an equity stake in Carbon Inc., which is a venture-backed startup company that owns related US Patent 9,216,546, US Patent 9,360,757 and others. LT reports a relationship with Digital Light Innovations that includes: employment, equity or stocks, and non-financial support. Methods for multi-material iCLIP are the subject of a pending patent application. The authors declare no other competing interests.Data and materials availability: We report that at the time of submission to the Proceedings, in August of 2022, part of the work presented here is accepted for publication in the journal Science Advances.1978References[1] J. Huang, Q. Qin, J. Wang. “A Review of Stereolithography: Processes and Systems” Processes.

gital Micromirror Device Based Microstereolithography for Micro Structures of Transparent Photopolymer and NanocompositesGeorge W. Hadipoespito, Yong Yang, Hongseok Choi, Guoqing Ning, and Xiaochun Li Department of Mechanical Engineering University of Wisconsin-Madison Madison, WI 53706 Reviewed, accepted August 19, 2003 Abstract: This paper describes a Digital Micromirror Device (DMD) based ultraviolet (UV) microstereolithography (µ-SL) system developed for rapid prototyping and manufacturing of micro 3D structures. Characterization experiments show that the developed the DMD-based imaging system irradiates an entire photopolymer layer at once, providing reasonable curing speed and good resolution at a low cost. 2D and 3D micro parts were fabricated. High frequency ultrasonic vibration (above 20 kHz) was experimented and verified that it can be used to significantly decrease the leveling time of viscous photopolymer.

High frequency ultrasonic vibration (above 20 kHz) was experimented and verified that it can be used to significantly decrease the leveling time of viscous photopolymer. Furthermore, micro parts were also fabricated in nanocomposites, which were obtained by ultrasonic mixing of the transparent photopolymer and nano-sized ceramic particles. High quality micro models fabricated by this novel process could be used for micro scale investment casting, tooling, devices, and medical applications.1. IntroductionMiniaturization is important in fabrication and commercialization of new industrial products that encompass complex geometries, high aspect ratios and complex microstructures. These miniature products can deliver a new generation of functionality and performance in a smaller volume.

These miniature products can deliver a new generation of functionality and performance in a smaller volume. Smaller-size cellular phones and notebook computers, hearing aids, nozzles for ink-jet printers, and accelerometers for car air-bag system are some samples of products made possible by rapid development in micro fabrication. Beside the commercial/consumer products and applications, 3D microparts are also of great interest from research fields in microrobotics, micromechanics, and microfluidics.Silicon-based microelectromechanical system (MEMS) techniques are very popular for the manufacture of micro sensors, electronics, actuators and other components in integrated microsystems. However, these MEMS techniques are two-dimensional (2D) processes involving multiple steps and requiring complicated processing procedures in a cleanroom environment.

However, these MEMS techniques are two-dimensional (2D) processes involving multiple steps and requiring complicated processing procedures in a cleanroom environment. These 2D techniques are normally difficult, if not impossible, to fabricate arbitrary shape and composition without the use of microassembly. The investment costs of these processes are expensive, and the choice of material is somewhat limited. To name a few of those processes are surface and bulk micromachining, LIGA (X-Ray Lithography, Electroplating, and Molding), and RIE (Reactive Ion Etching) [Madou 2002].Recently, microstereolithography (µ-SL) process emerged as a freeform process capable of creating 3D structures with complex geometries. Microstereolithography has its root from the stereolithography process rather than integrated circuit (IC) technology. Stereolithography is a process of creating 3D solid models in a layer-by-layer fashion from a 3D CAD model using13photopolymer as a based material.

Stereolithography is a process of creating 3D solid models in a layer-by-layer fashion from a 3D CAD model using13photopolymer as a based material. A focused ultraviolet (UV) laser, coupled with a galvano- mirror, is used to cure photopolymer in a vector-by-vector manner [Jacobs 1996, Beaman 1997, and Prinz 1997].The core idea of microstereolithography processes is similar. A solid 3D model from any modeling software (Pro/E, Unigraphics, SolidWorks, etc) is tessellated into array of triangles, and output in STL (SteroLithography) file format. Using slicing software, this tessellated model is then sliced into layers according to the predetermined thickness. The software then creates some output files that will be read and used directly to build the part. The output build files are then used to build the part on the microstereolithography apparatus.

The software then creates some output files that will be read and used directly to build the part. The output build files are then used to build the part on the microstereolithography apparatus. The microstereolithography processes can be classified into two main categories, based on how each layer is build [Bertsch 1999 and 2000]:1. Vector-by-vector process: a layer is formed by raster scan of a laser beam. The laser beam is stationary and focused precisely to a very fine spot on the surface of the photopolymer. The container and the material are moved horizontally in order to build the layer. Integral process: each layer is cured by one irradiation only. A dynamic mask/spatial light modulator is used. For this type of process, the slicing software gives output in image files (usually monochrome bitmap files); each file contain the cross sectional image of one single layer. These files are then used to drive the dynamic mask.

These files are then used to drive the dynamic mask. After a layer is formed, a microstage holding the part is dipped into the photopolymer vat, allowing fresh layer of resin to coat the cured layer. Then the microstage is raised again, leaving only a thin layer of liquid photopolymer (usually in the order of several microns) between the cured layer and the surface of the polymer vat. The fresh layer is then cured, and then the recoating process repeats. Once building process is done, the part is washed with the appropriate solvent to remove the uncured polymer.2.The integral stereolithography process is much faster than the vector-by-vector process. By employing a dynamic mask, e.g. Liquid Crystal Device (LCD), this process presents a technical breakthrough in the field of rapid prototyping.

By employing a dynamic mask, e.g. Liquid Crystal Device (LCD), this process presents a technical breakthrough in the field of rapid prototyping. As an emerging dynamic mask, the performance of Digital Micromirror Device (DMD) is better than that of the LCD since DMD’s fill factor of each pixel (85%) and its reflectivity (71%) are much better than those of the LCD (64% fill factor and 21% light transmission) [CRL Opto, DLP Technology]. In other words, when a pixel is in “active” or “on” state, the DMD can pass light more efficiently than the LCD. LCD-based µ-SL technology has been extensively studied [Bertsch 1999 and 2000]. EPFL (Swiss Federal Institute of Technology) at Lausanne, Switzerland, University of Sussex in United Kingdom, INPL (Lorraine National Polytechnic Institute) in France, and Saitama University in Japan are to name a few. Long layer settling time remains a difficult issue.

Long layer settling time remains a difficult issue. Research team from Saitama University in Japan used DMD-based µ-SL to build 3D parts [Kaneko 2001]. A vibration generator working at frequency of 50 Hz was sued to reduce the leveling time and enhance the z-resolution by reducing layer thickness [Akahane 2001]. However, its layer thickness was still at approximately 100 µm. It should also be noted that Stanford Research Institute (SRI) has developed a projector- based imager system for ceramic-polymer prototyping [SRI 1999]. However, the system uses visible light, which is prone to interference and the system developed is suitable only for macroscale applications.142. Experiment ApparatusA microstereolithography (µ-SL) apparatus has been developed, as shown in Figure 1. The system employs a Digital Micromirror Device™ (DMD) as a dynamic mask/spatial light modulator, which eliminates the need to produce multiple masks to cure different layers of polymer.

Optical SystemThe light path on the apparatus begins from the UV light source, identified by number 1 in Figure 1. The light is then transmitted using a UV light guide (number 2 in Figure 1) into a modified DMD-based multimedia projector (BenQ DX550). The light path starts at the light pipe (number 3 in Figure 1) and ends at the TIR (total internal reflection) prism pair (number 7 in Figure 1). A set of focusing lenses, identified by number 8 on Figure 1, is placed outside of projector to focus the reflected image from DMD onto the surface of polymer bath in a suitable size.To supply UV light, a broadband GreenSpot™ UV Spot Curing System from UV Source, Inc. was used. The UV Spot Curing System consists of a high-pressure mercury lamp, a 100W power supply and ballast system, and a programmable shutter. The shutter is electronically controlled, and can be adjusted in increments of 0.01 second up to 99.99 seconds.

The shutter is electronically controlled, and can be adjusted in increments of 0.01 second up to 99.99 seconds. The spectrum of the broadband UV light ranges from 300 and 470 nm, and has a peak at 365 nm. This external UV light source was selected to avoid difficult modifications required to integrate the stand- alone high-pressure mercury bulb into the projector. The light guide assembly was also supplied by UV Source, Inc. The light guide is liquid-filled, and provides good UV transmission between 300 and 500 nm.15The DMD has a resolution of 1024x768. It provides rapid change of image patterns (up to 10,000 paterns/s). It consists of thousands of micro mirrors, each with a size of 13.7 µm x 13.7 µm. Each individual mirror element can be rapidly electronically switched (up to 10,000 Hz) between “on” and “off” state. The light is thus selectively reflected, according to a designed image pattern, out of the projector into the focusing lens.

The light is thus selectively reflected, according to a designed image pattern, out of the projector into the focusing lens. For this study, the lens system was designed to provide approximately 1:1 ratio between the DMD pattern and the focused image. Silica lens was used since silica has good transmission in the UV region.2.2. PhotopolymerSuitable photopolymers are critical for the µ-SL process. Important characteristics, such as viscosity, sensitivity, and wavelength response need to be considered. After UV curing, polymers will also exhibit different thermal, mechanical, and optical properties.Several photopolymers from companies that specialize on stereolithography were evaluated. The DSM Somos 10120 WaterClear™ was selected because of its low viscosity (~130 cps @ 30º C), its relatively low critical exposure (Ec = 9.7 mJ/cm2), and its transparency in both liquid and cured state.

ln(E) curve)Index of Refraction, nE10 (Exposure that gives 0.254 mm (.010 in.) thickness) Table 1. Properties of DSM Somos WaterClear™ 10120 photopolymer. [Data from DSM Somos]1.51 812.3. Motion ControlTo provide an accurate linear movement of the building platform (number 11 on Figure 1), a PC-controlled three-axis micro-stage (number 10 on Figure 1) from Anorad Inc. was implemented. Two perpendicularly stacked LW-7 stages provide lateral (x and y directions) movements, while an Anoride 7-4 stage enables the platform to move vertically (in z direction). The stage system has a resolution of 30 nm and can travel in speed up to 200 mm/s.2.4. SoftwareTwo different types of software were used for the apparatus: CNC-2000, and Microsoft PowerPoint.

The stage system has a resolution of 30 nm and can travel in speed up to 200 mm/s.2.4. SoftwareTwo different types of software were used for the apparatus: CNC-2000, and Microsoft PowerPoint. The CNC-2000 controls the motion of the stages; manual input of a single command line is possible, as well as automatic input of a series of commands in G-code format.16Microsoft PowerPoint was used to create monochromatic patterns and images and transferring those to the DMD. Further software development is needed to automate the solid model slicing and image generating.3. Experiment procedureThe experiments were performed to characterize the apparatus and evaluate its capability for fabrication of meso and micro-sized 3D objects. The results will be used in selecting important variables and optimizing the fabrication process.3.1.

The results will be used in selecting important variables and optimizing the fabrication process.3.1. Experiment on curing depth and exposure time experimentThis series of experiments was carried out to better understand the relationship between the exposure time and cured layer thickness. The experiment was conducted by fully curing a support polymer layer on top of a silicon wafer, as shown in part 2 in Figure 2. A circular part on this layer was intentionally left uncured for top layer measurement purpose. Then a fresh layer of liquid polymer was recoated above the cured layer by dipping and raising the building platform. Finally, this top layer was exposed to a focused UV light within a predetermined period of time (part 3 in Figure 2). The cured thickness was then measured by use of a microscope, as illustrated in part 4 in Figure 2. The relationship between exposure time and cured layer thickness can be determined.Fig. 2.

The cured thickness was then measured by use of a microscope, as illustrated in part 4 in Figure 2. The relationship between exposure time and cured layer thickness can be determined.Fig. 2. Experiment on curing depth versus exposure time3.2. Experiment on settling timeThe apparatus builds microparts by stacking thin layers of photopolymers, preferably 5- micron thick for each layer. Conventional stereolithography apparatus normally utilizes a blade to achieve a desired layer thickness, normally 50~100 µm. However, this technique can not be applied on a µ-SL apparatus since the blade will destroy the micro part being built. A µ-SL apparatus usually relies on gravity to settle the polymer layer into a thin layer.The experiment was conducted by lowering the building platform, and then raising it within the polymer bath up to height that is 10.0 µm below the surface of the photopolymer.

Once the platform is raised, UV light was turned on periodically to cure a different small part of the settling polymer. The cured layer thicknesses were measured by a Tencor Alpha-Step 200 profilometer.3.3. Experiment on resolution17There are two different resolutions on a µ-SL apparatus: lateral resolution, and vertical resolution. Lateral resolution is decided by the size of the individual mirror on the DMD as well as the focusing lens system. Vertical resolution is simply determined by the layer thickness; the smaller the layer thickness, the better the vertical resolution. However, thinner layer thickness demands more layers to build a 3D part.The experiment was carried out to quantify the lateral resolution of the system. The experiment was conducted by curing a pattern of lines with different widths. After the sample was developed, the narrowest line width indicated the resolution of the apparatus.3.4.

The experiment was conducted by curing a pattern of lines with different widths. After the sample was developed, the narrowest line width indicated the resolution of the apparatus.3.4. Fabrication of Micro Parts2D and 3D patterns/parts were fabricated to assess the capability of the apparatus. A microgear was of interest because it is a mechanical component containing small features. A microgear with 1.5mm lateral diameter was designed to be built using the apparatus, as shown in Figure 3a. To increase the complexity of the micropart, a cylinder and a second, smaller gear with fewer teeth (Figure 3b) were to be built on top of the base microgear.A micropart with overhanging feature was also of particular interest. One attractive capability of the µ-SL process is that it is able to build overhanging features without the need of supporting structures. Hence, a micropillar with 1mm x 1mm lateral dimension with an overhanging section was designed, as shown in Figure 3c.Fig. 3.

Hence, a micropillar with 1mm x 1mm lateral dimension with an overhanging section was designed, as shown in Figure 3c.Fig. 3. Shown above are the cross-section profile of the base gear (a, left) and that of the second, smaller gear (b, middle) (the pictures are not in the correct scale). A CAD model of a micropillar with overhanging section was to be built using the apparatus (c, right).3.5. Experiment on ultrasonic-assisted layer settlingHigh frequency ultrasonic vibration (above 20 kHz) was experimented in order to significantly decrease the leveling time of photopolymer. As shown in Figure 4, the experiment setup includes a 20 KHz ultrasonic ultrasonic transducer, a vibration table, and a silicon wafer. The thickness of the droplet of photopolymer on top of the silicon wafer was measured against time. A droplet of photopolymer of 14.6 µl was placed on top the silicon wafer under ultrasonic vibration.

A droplet of photopolymer of 14.6 µl was placed on top the silicon wafer under ultrasonic vibration. By measuring the diameters of settling droplets, the average thicknesses of the settling droplet layers were determined.18SiliconPolymerUltrasonic Transducer (20 KHz)Vibration tableFig. 4. Diagram of ultrasonic-assisted layer settling3.6. Nano-composites preparationThe experiment was carried out to determine the feasibility of mixing polymer with nano- sized ceramic particle. Polymer-nano-composite material was of interest because it had the potential of improving mechanical properties of the fabricated parts.High intensity ultrasonic waves were used to mix nano-sized Si3N4 (30 nm) with the photopolymer. A viscometer was used to measure the viscosity of the resulting polymer-nano- composite mixtures. The measurement result was plotted and used to determine the suitable nano-composite content in the mixture.4. Results & Discussions4.1.

The measurement result was plotted and used to determine the suitable nano-composite content in the mixture.4. Results & Discussions4.1. Curing depth and exposure timeFigure 5 shows the result of the curing depth versus exposure time. The linear fit of the data points on the plot suggests the cured layer thickness is proportional to the exposure time. The plot suggests that the polymer would not start to cure if the exposure time is less than 1.67 seconds in this µ-SL system. Other experiments verified this behavior. It might be attributed to the use of broadband UV light source and the fact that the photopolymer used is optimized for curing at 355nm wavelength. It might be better to use low power UV laser at 355nm to obtain optimized performance for this particular polymer. However, currently these UV lasers are expensive.

It might be better to use low power UV laser at 355nm to obtain optimized performance for this particular polymer. However, currently these UV lasers are expensive. One possible solution would be to find or develop photopolymers that have better response curve with respect to the spectrum of the broadband UV light.1925002000y = 162.49x - 270.87 R2 = 0.9872m ÷, h t p e d1500d e r u C1000500002468101214Exposure time, secondsFig. 5. Cured depth (cured layer thickness) versus exposure time on DSM Somos WaterClear 10120.4.2. Settling timeFigured 6 shows how the polymer surface settled naturally (decreased in thickness with respect to the building platform) on top of silicon wafer as time elapsed. The plot suggests that the settling polymer layer thickness is proportional to the elapsed time. However, the settling time of the polymer was surprisingly long to reach 100 µm.

The plot suggests that the settling polymer layer thickness is proportional to the elapsed time. However, the settling time of the polymer was surprisingly long to reach 100 µm. The result suggests that 895.36 seconds would be necessary to obtain a polymer layer of 10 µm thick. Such a long settling time is not satisfactory for a rapid prototyping system. Further study is needed to find a more suitable photopolymer or to develop other means to spread thin polymer layer in this µ-SL system.To overcome this settling problem, ultrasonic vibrations could help to achieve a significantly shorter layer settling time. Preliminary settling time experiments with ultrasonic vibrations were conducted, and results are presented in the subsection 4.5.180]m µ[160140120y = -0.1681x + 160.51 R2 = 0.9698h t p e dr e m y o Pl1008060402000100200300400500600Elapse time [sec]Fig. 6.

6. The height of the polymer surface measured from the building platform versus elapsed time.204.3. ResolutionBy comparing the line widths on the computer screen to the cured line widths, the image ratio can be determined. A ratio of 1:20.3 can be achieved between the line width on the screen and the actual cured line width.By use of microscope and SEM, the lateral resolution of the apparatus is determined to beapproximately 20 µm. It dictates the smallest feature possible in X-Y plane to be built.Fig. 7. Shown above are a microscope capture and SEM pictures of lines and grids patterns generated using µ-SL apparatus.4.4. Fabrication of 3D Micro PartsFigure 8 shows SEM pictures of 3D microparts fabricated with the µ-SL apparatus. Individual tooth can be distinguished on both micro-gears; however it seems some features were over cured.

Individual tooth can be distinguished on both micro-gears; however it seems some features were over cured. It is believed that optimization of process parameters will be necessary to solve this problem.The last image in Figure 8 shows the micro-pillar. It seems no obvious overhanging feature appears. It is likely that the overhanging feature was washed out, possibly due to the recoating process. Further experimental study and process optimization are being carried out to obtain more sophisticated parts.Fig. 8. SEM pictures of 3D microgears, and a micropillar (right)214.5. Ultrasonic-assisted Layer SettlingComparison between the results with and without ultrasonic vibration indicates that ultrasonic-assisted settling is very effective.

Ultrasonic-assisted Layer SettlingComparison between the results with and without ultrasonic vibration indicates that ultrasonic-assisted settling is very effective. It shines light for the future implementation of ultrasonic-assisted layer settling mechanism in the developed DMD-based µ-SL system.300r e t e m o r c i m250200(150s s e n k c i h T10050without ultrasonicwith ultrasonic002040 Time (s)6080Fig.9. Polymer settling time on top of silicon wafer4.6. Polymer Nanocomposite PartsFigure 10 shows the viscosity trend of the photopolymer matrix nanocomposite as thenanoparticles content is increased. Based on the result, the nanocomposites having 1.0 weight % nanoparticles might have the combination of good strength improvement and their viscosity still remains low enough for recoating process. Thus photopolymer was mixed with 1.0 weight % by use of ultrasonic mixing.

Thus photopolymer was mixed with 1.0 weight % by use of ultrasonic mixing. The resulting mixture was then used to fabricate a micro-gear shown in Figure 11. Visual observation under a microscope suggests that the nanoparticles within the microgear were dispersed quite uniform.500045004000y = 210.2e1.4583x R2 = 0.9915] p c [35003000y t i s o c s i V2500200015001000500000.511.522.5Weight of nano particles in mixture [%]Fig. 10. Viscosity of photopolymer-nanoparticle mixture (measured at 1.0 rpm)22Microgear of nanocompositeFig. 11. A microscope picture of a microgear made of polymer-nano- composite.5. ConclusionA DMD-based microstereolithography apparatus has been realized. The apparatus attempts to combine the versatility of rapid prototyping, and the capability of fabricating high resolution micro structures. 2D and 3D microparts with 20 µm lateral resolution are realizable.

2D and 3D microparts with 20 µm lateral resolution are realizable. Characterization experiments show that the developed the DMD-based imaging system irradiates an entire photopolymer layer at once, providing reasonable curing speed and good resolution at a low cost. Further study and process optimization will be needed to improve the quality of the micro parts.High frequency ultrasonic vibration was experimented and verified that it can be used to significantly decrease the leveling time of viscous photopolymer. Uniform photopolymer matrix nanocomposites were obtained by ultrasonic mixing. The nanocomposites promise to improve properties of fabricated parts. Micro parts were also fabricated in nanocomposites.6. AcknowledgementsThe authors would like to thank Mr. Zhiwei Li for his assistance in the SEM pictures. We are grateful for the financial support from the Wisconsin Alumni Research Foundation (WARF).237.

299-304, 2001.Prinz, F. B., chair, "Process Overviews" in “Rapid Prototyping in Europe and Japan”, Vol. I,Analytical Chapters, International Technology Research Institute, Loyola College in Maryland, March 1997.Provin, C., and Monneret, S., “Complex Ceramic-Polymer Composite Microparts Made byMicrostereolithography,” IEEE Transactions on Electronics Packaging Manufacturing, Vol. 25, n1, pp. 59-63, January 2002."SRI's Direct Photo Shaping (DPS)", Rapid Prototyping Report, pp.6-7, July 1999. UV Source, Inc. (http://www.uvsource.com/home.html)24.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDIGITAL MODEL GENERATION STRATEGY FOR PRINTING MINIATURIZED REPLICA OF HISTORICAL CAST IRON OBJECTS – A USE CASE STUDYD.Borstell*, B.Friedhofen**, D.Kunz*Research Laboratory for Additive Manufacturing, Koblenz University of Applied Sciences, 56075 Koblenz, Germany ** Cast Iron Museum, Stiftung Sayner Hütte, 56170 Bendorf-Sayn, GermanyAbstractThe Cast Iron Museum at Sayn (Germany) hosts a wide range of historic cast iron objects ranging from highly detailed fly-size figures to oversize eagles and church bells. Replicating the historical cast iron, i.e. free formed objects, as miniaturized 3D prints for a new exhibition requires scanning to create the digital model.High quality of the digital model is a fundamental prerequisite for the production of the miniaturized 3D print.

The influence of software selection and settings during data preparation from point cloud, mesh and finally solid is highlighted. Printing process selection is the final step in achieving the required print quality.A scanning and model generation strategy is developed and described using several exemplary objects from the Sayn Museum. The final print quality resulting from the selected scanning process is evaluated.Motivation and Object SelectionThe Iron Foundry at Sayn (Germany) dates back to 1769. In 1815 the Prussian State and later in 1865 Alfred Krupp1 took control. Various cast iron goods have been produced ranging from industrial applications (guns, church bells, etc.) to household products and fine arts goods.

25 mm), (c) Bust of “Elisabeth“ (height approx. 35 mm), (d) Bust of “Benedicta von Löwendal“ (height approx. 800 mm)efghFigure 2: (e) Garden Chair by Karl Friederich Schinkel [4], (f) Vase (diameter 740 mm, height 900 mm), (g) Church Bell with inscription “Sayner Hütte“ and ornament of Prussian Eagle (bottom diameter 1250 mm, height 1200 mm), (h) Prussian Eagle (wingspan 1050 mm)The foundry was closed in 1926, but after the restoration started in 2004 a Cast Iron Museum was established hosting a wide range of original objects. For 2023 a new exhibition was planned “3D Printing Meets Iron Art Casting” for which miniaturized replica of these objects were needed. Except for the “Sayn Fly”2 all listed objects were scanned and printed.

Except for the “Sayn Fly”2 all listed objects were scanned and printed. Practical experiences from scanning the bust of “Benedicta von Löwendal” (d), Chair by Schinkel3 (e), vase (f) church-bell (g) and eagle (h) result in case-specific strategic model generation rules.State of the ArtA strategy is a plan of action designed to achieve a long-term or overall aim. Scanning free formed objects with the task to produce miniaturized prints for an exhibition requires such a plan. The simple plan “scan – slice – print“ has to be detailed and fundamentally improved to lead to success.A digital model is the absolute prerequisite for manufacturing objects additively.

The simple plan “scan – slice – print“ has to be detailed and fundamentally improved to lead to success.A digital model is the absolute prerequisite for manufacturing objects additively. The quality of the digital model is therefore decisive for the quality of the printed part and its creation is part of the Additive Manufacturing Product Development Process (AMPD process).2 “Sayner Mücke“, cast from only one drop of molten iron, weight 0,27 gr., legs were made from steel wire 3 Karl Friedrich Schinkel (1781 – 1841), Prussian building official, master builder, architect and designer1802dThe current AMPD process consists of four steps: product planning, product design, work preparation and product manufacturing [1].According to Auth [1], work preparation covers the pre-process of AM whereas product manufacturing covers the in-process and the post-process of AM.

For designing a new technical object, a modelling process is used by 3D CAD modelling with an optional topology optimization. The created volumes and shapes are based on simple (primitive) analytically defined, i.e. exact, geometries. Topology optimization may create free formed surfaces defined by splines. The digital model is defined during the product design phase. Replicating does not require a planning and a design phase, since the object already exists.Figure 4: AM process and modified AM processReplicating existing objects with free form surfaces requires a scanning process. Scanning produces a point cloud which is transformed into a surface model by meshing creating an STL-file. The solid is then the base for slicing and printing. The detailed description of the replication process will identify the individual steps to be taken and highlight parameter settings and methods leading to an optimal digital model.

The detailed description of the replication process will identify the individual steps to be taken and highlight parameter settings and methods leading to an optimal digital model. The final quality depends on parameters within the process which can be influenced only within certain limits. The steps of the replication process and their significance to the final print quality are highlighted and discussed. The complete replication process has two steps: scanning and modifying. The modification step may be detailed into repairing and refining. Repairs are modifications to theFigure 5: Replication process1803model which are necessary to print the model. Refinements are modifications bringing the surface geometry as close to the original as possible.

Refinements are modifications bringing the surface geometry as close to the original as possible. However since no original digital model exists, no quantitative measurement of the model´s deviation from the original is possible.Slicing and printing are part of the replication, but the digital model is fully defined afterall modificationsScan ProcessTo execute the scanning process hard- and software are important items which have been selected beforehand. Object size and scanner resolution are the leading parameters. For scanning the bust of “Benedicta”, chair, vase and eagle a “Shining EinStar 3D” handheld scanner is used. It is connected by cable to a “HP ZBook vPRO” with an Intel Core i7 processor and two graphics cards (one Nvidia Quadro M100M and one Intel HD Graphics 530) to allow fast processing of large amounts of data.

The ZBook also hosts the Shining proprietary scanning software “ExStar” which controls the scanning process and allows displaying the point cloud and meshing the point cloud into an STL-format. Modifications will be done in “ExStar”, Autodesk “MeshMixer“, Artec “GOM Inspect“, “Fusion 360“ as well as Dassault Systèmes‘ “Solid Works“.The “Shining EinStar 3D“ is a structured-light handheld scanner equipped with two stereo cameras, one RGB camera and an infra-red VCSEL Laser (Vertical Cavity Surface Emitting Laser) structured light source. The object is fixed and the scanner is moved around the object making it suitable for large static objects. Resolution will reach 0,1 mm point distance. Structured-light scanners project a precisely calibrated pattern (parallel lines, stripes or a grid) onto the surface of the object. The corresponding pattern is distorted when it hits the object.

Structured-light scanners project a precisely calibrated pattern (parallel lines, stripes or a grid) onto the surface of the object. The corresponding pattern is distorted when it hits the object. The pattern is captured by the scanner's built-in stereo cameras in a high-frequency scan and processed by the integrated scanning software through active triangulation. Due to the defined positions of light projector and cameras, the coordinates of each single surface point can be precisely calculated. Data of the RGB camera allow determination of each point´s color. Structured light scanning allows objects to be captured in the sub-millimeter range without touching them. [2]Figure 6: “Shining EinStar 3D“ Scanner (left [5], middle), distorted light pattern (right [2])Result of the scanning process is an ASCII file containing the coordinates of each point of the point cloud which is displayed in the “ExStar” software. Each single point of the cloud is visible.

Each single point of the cloud is visible. The point cloud is an incomplete representation of the physical object due to scanning faults, missing surface between the points and the non-solid character. Modification of the model by manipulating single points in the ASCII file would be tedious and ineffective.1804However, it is the starting point for all further modifications leading to the optimal digital model with the least possible deviations from the original object.Figure 7: Point cloud - view of the back of the chair from above and extract from ASCII fileCast iron surfaces are rough. When scanning large areas, e.g. the bell or the vase operators tend to scan the same area multiple times to be sure to catch every detail of the surface. Each time different points will be scanned. As a result, the point cloud files become huge.

Each time different points will be scanned. As a result, the point cloud files become huge. File sizes of several GBs were experienced.Figure 8: Typical cast-iron surface of the bell (left,), optical feedback when scanning the cast iron surface of the bell (right), [14]1805Point Cloud ModificationThe scanning step has been completed with the creation of the point cloud, all further repairs and refinements are listed below. The necessity of repairs becomes evident by the integration of some automatic modifying functions in the scanning software “ExStar”.

The necessity of repairs becomes evident by the integration of some automatic modifying functions in the scanning software “ExStar”. The following table highlights the possible combinations of modifications and the manual respectively automatic routines provided by the available software packages:ModificationExStarMesh MixerGOM InspectSolid WorksTransformation into meshed surface (STL)automaticXClosing of holes / gapsRemoval of floating partsSmoothing of surfaceautomatic manual automatic manual automatic manualXXXX XXXAdding analytically defined geometry Replacing Elements Define origin Align with horizontal planemanualmanual manual manualXXXX XXTo allow easy modification the cloud is transformed into a meshed surface.

Theindividual points of the cloud are connected to form triangles:Figure 9: STL of the eagle (left) enlargement of the triangles forming the surface (right)The triangle mesh only approximates the real surface of the object, a deviation from the original is evident. A large number of triangles will result in a more detailed surface model (closer to the original) and is desirable. “ExStar” allows a maximum number of 20 million triangles in the mesh, leading to large STL files and meshing times between a few minutes and more than 24 hours depending on the size of the input file. The number of triangles can only be selected at this point and will have to be considered carefully. Two small exemplary point cloud files were meshed with all parameters set to minimum values and results noted:1806Figure 10: Exemplary point cloud files, 8276 KB (left: small) 53913 KB (right: large)Size of input point cloud file 8276 KB 8276 KB 8276 KB 53913 KB 53913 KB 53913 KBMax.

number of triangles 20.000 200.000 2.000.000 20.000 200.000 2.000.000Size of meshed file 974 K 9759 K 187960 K 981 K 9763 K 97654 KMeshing time 60 sec 105 sec 140 sec 100 sec 125 sec 235 secThe results are as expected: meshing times and STL-file sizes rise with original file size and maximum number of triangles. Therefore, large input files will increase meshing times and STL-file sizes.Closing of holes and gaps, floating parts removal and smoothing of the surface are integrated automatic functions of the “ExStar” meshing routine. Gaps and holes are areas where the distance between points of the cloud becomes too large to be connected to the next point. These can be detected and closed automatically but the large point distance will result in an imprecise approximation. “ExStar” offers three basic meshing qualities (“Non-Watertight, Semi-Watertight and Watertight”) and a range of optimization options.

“ExStar” offers three basic meshing qualities (“Non-Watertight, Semi-Watertight and Watertight”) and a range of optimization options. These options include smoothing of the mesh, removal of floating parts, filling of small holes and the removal of spikes. Filling of large gaps or holes may exceed manual skills. However, the final consequences of an automatic repair of such a point cloud fault becomes evident in the model and the printed eagle. Adding more points should be seriously considered. However, the scannerFigure 11: Gap in the point cloud (left), resulting visible structures of auto repair routines in slicer (middle) and printed object (right)software mostly lacks efficient functions to add new points to an existing point cloud and the scan may have to be repeated in total.1807Floating parts develop out of small point clouds which cannot be associated to the main object cloud. Sometimes they are caused by background points which have been caught in the scan.

Sometimes they are caused by background points which have been caught in the scan. These parts are obsolete and must be removed. After meshing the digital object, the mostly automatic process will have created a surface model (STL-file) cured of many faults.Figure 12: Floating parts in the point cloud of the eagle, “ExStar“ meshing optionThe semi-automatic process provides a welcome assistance to the modification process. The algorithm identifies “standard” faults and executes the repairs. Parameters of identification and repair limits can be set. It is up to the user to decide about their optimal setting.After meshing the digital object is still an incomplete surface model and further improvements are required. The Autodesk “MeshMixer” is a freely available software providing a range of functions to improve model quality. The display will identify open areas by highlighting outside surfaces in white, inside surfaces in pink and rendering the edges in blue.

The display will identify open areas by highlighting outside surfaces in white, inside surfaces in pink and rendering the edges in blue. Holes and gaps become clearly visible. The “Inspector” function will highlight all identified holes and offer an automatic “Hole fill” algorithm. The gaps and holes marked in red and magenta exceed the capabilities of the algorithm and must be cured manually.Figure 13: “MeshMixer“ vase with gaps (left), “Inspector“ mark up (middle), vase after automatic hole filling (right)Although tempting, not all holes should be closed automatically. The historic original object (vase) will contain some cracks, gaps or bumps in the original structure which need to be kept and require careful manual modifications.

The historic original object (vase) will contain some cracks, gaps or bumps in the original structure which need to be kept and require careful manual modifications. Replication of the original geometry is the goal, not the perfectly approximated surface.1808Figure 14: Manual modification of vase details representing the original (faulty) structure: (a) original vase, (b) meshed surface in “MeshMixer“, (c) manually closed holes [12]Smoothing the surface is part of the “ExStar” meshing routine. However only threedifferentiation possibilities in the parameter settings (“low, med, high”) let us expect an unwanted change in the original geometry. “MeshMixer” offers a range of “sculpting” functions which will also be helpful to smooth specific areas manually. Additional smoothing was not applied to the eagle (Figure 11).

“MeshMixer” offers a range of “sculpting” functions which will also be helpful to smooth specific areas manually. Additional smoothing was not applied to the eagle (Figure 11). Relying on the automatic hole closing routine has proven to be deceptive.Adding analytically defined geometry is possible in “MeshMixer” by adding planes and filling the resulting voids with matter. The geometry of the base of “Benedicta’s” bust is cubically shaped but has not been completely captured in the scan because the object was standing on a table.Figure 15: Filled up base of “Benedicta“ [14]1809Alternatively, planes can be used to cut off surplus matter. In case of the bell a horizontalplane is added 5 mm above the lower edge in order to create a straight lower brim.Figure 16: Bell mesh before cutting (left), bell with cutting plane in “GOM Inspect“, bell after cutting base [14]Replacing scanned elements by mirroring at a plane of symmetry was used to create the model of the chair.

The rod structure of the chair made it extrelemy difficult to capture the single rods without scanning structural elements in the background at the same time.Figure 17: Incomplete preview, complete half (point cloud), joining mirrored halfs in “Mesh Mixer”Replacing scanned elements by analytically modelled elements was used to cure problems with the bust of “Benedicta” and the vase. Both could have been repaired in “MeshMixer”, however using analytically modelled elements was easier. The scanner could not reach all the way into the inside cavity of the vase where a huge gap in the inside surface was left after meshing. The concerned area of the vase was at the inside diameter and therefore rotationally symmetrical. Analytically defined surfaces form the basis of modelling in CAD systems. Exporting the STL-file to “SolidWorks”, sketching the inside onto the scanned contour and creating the new inside surface by using a “revolve base/boss” feature proved to be effective.

Exporting the STL-file to “SolidWorks”, sketching the inside onto the scanned contour and creating the new inside surface by using a “revolve base/boss” feature proved to be effective. However, the deviation from the original inside surface is significant.1810Figure 18: Inside after meshing (left), replacement sketch in “SolidWorks“ [12] (middle), repaired inside surface [12] (right)The bust of “Benedicta” features an inscription at the front of the base which proved to be difficult to scan by hand.

The inscription was deleted in “MeshMixer”, modelled in a “Fusion 360” and inserted into the model again.Figure 19: Original inscription (top left), scanned inscription (top middle [14]), modelled inscription (top right [14]), base with deleted inscription (bottom left [14]), base with modelled inscription added (bottom right [14])Aligning with horizontal plane and the definition of the origin by aligning the object with the world coordinate system are the last steps to complete the digital model.

“GOM Inspect” provides effective functions to define the object´s origin by placing planes on the base surfaces, aligning straight lines with the edges and locating the origin at their intersection.Figure 20: Base of the eagle with reference planes in green and object´s origin (yellow arrow) [11]1811Aligning the object´s coordinate system with the world coordinate system is done by referencing the position with the grid display in “GOM Inspect”:Figure 21: Aligning “Benedicta“ with the world coordinate system [14]PrintingResin color options and surface quality led to the selection of stereolithography as printing process. Bell, eagle, “Benedicta” and chair were printed on a Stratasys NEO 450 printer. Size was scaled down to approximately 25%.

Bell, eagle, “Benedicta” and chair were printed on a Stratasys NEO 450 printer. Size was scaled down to approximately 25%. Miniaturization was necessary due to the build space restrictions, but also reduces the level of details and consequently approximations loose significance.Figure 22: Bell, eagle, “Benedicta“ and Schinkel Chair printed on a Stratasys NEO 4501812SummaryThe process “scan – slice – print” starts with the selection of the appropriate soft- and hardware. Software selection must consider that specific model modification functionalities may be required which exceed the capabilities of the proprietary scanner control software. Research about available software, required additional capabilities and their specific application in the context of replicating historical (i.e. “damaged” objects) seems trivial but is extremely important.Object size in this case required the use of handheld scanners which introduces the “human factor” into the process.

“damaged” objects) seems trivial but is extremely important.Object size in this case required the use of handheld scanners which introduces the “human factor” into the process. The rough but otherwise monotonous surface of the bell allowed good manual capture by the scanner, however leading to multiple scans. Multiple scanning of the same surface area led to huge file sizes and long scanning times. The rod structure of the chair could not be fully captured manually and needed extensive manipulation of the model afterwards.After scanning, necessary and optional modifications and the limitations of automatic routines have to be identified. A careful selection between automatic and manual processes is expedient as well as definition of parameter settings for all transformations (e.g. optimize setting of number of triangles for meshing). Slicing software selection was found to be no issue. The proprietary slicer-software of the printers was used which included all necessary functions.

Slicing software selection was found to be no issue. The proprietary slicer-software of the printers was used which included all necessary functions. Although the digital model is completed before printing, the printing process will influence the data modification process and may be considered. The surface quality and the color range of the stereolithography-process proved to produce the highest quality prints and therefore demands a matching model quality. Printing at 25% size was advantageous.Even if all parameters are set to the optimum values during the replication process, the digital model will only approximate the original object and a modelling process will always be more exact. Lack of original digital object data however limits any statement about the quality of the model to subjective statements.References[1] C. Auth, D. Borstell and R.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA digitally-driven Hybrid Manufacturing process for the flexible production of engineering ceramic componentsJ. Hinton1*, D, Basu, D. F. Flynn2, R. A. Harris1, R. W. Kay11.Future Manufacturing Processes, Institute of Design, Robotics and Optimisation Engineering,University of Leeds, West Yorkshire, LS2 9JT, UK2.Herriot-Watt University, Edinburgh Campus, Edinburgh EH14 4AS, UKAbstractCeramic materials are a versatile class of materials with numerous applications across a range of industrial sectors. Predominant methods of manufacturing ceramic components use template-driven methods, which hampers responsiveness and impose significant design constraints. This has driven significant interest towards digitally-driven manufacturing approaches, primarily, additive manufacturing.

This has driven significant interest towards digitally-driven manufacturing approaches, primarily, additive manufacturing. Additive manufacturing has demonstrated the rapid production of bespoke and highly complex geometries and designs direct from digital data without the need for component specific tooling. Yet, when used in isolation these techniques are restricted by uncontrollable porosity, high shrinkages during firing plus a lack of process-compatible materials. This paper presents the research and development of a new hybrid manufacturing process chain for the agile production of engineering grade ceramics components. The combination of high viscosity ceramic paste extrusion, sacrificial support deposition and subtractive micro-machining has yielded complex monolithic ceramic components with feature sizes of 100µm, part densities of ~99.7%, surface roughness down to ~1µm Ra and 3-point bend strength of 218MPa.

Since a wide range of materials can be formulated into visco-elastic pastes they can be readily deposited using this approach.IntroductionEngineering ceramics are broadly defined as inorganic, non-metallic materials [1] and are extensively used across broad and varied spectrum of applications and industries including; healthcare [2], oil and gas [3], power generation, automotive [4] and aerospace [5]. Ceramics exhibit a range of desirable material characteristics including high hardness, excellent thermal and electrical insulation, high resistance to wear, erosion and corrosion [6]. However, these characteristics present a number of challenges during subsequent formative and manufacturing processes.Conventional methods of manufacturing precision-engineered ceramic components often involve multiple formative and finishing processes that are segregated by a profiled thermal cycle.

The hard and brittle nature of ceramic materials results in bulk ceramic being unsuitable for formative and large-scale material removal processes. Therefore, ceramic powder with a narrow particle-size distribution, is dispersed within an organic binder and additional additives to improve processability and component performance. The subsequent matrix exhibits suitable fluidity and characteristics to undergo formative processing such as injection moulding [7], extrusion [8] and machining [9][10]; yielding a near-net-shape green state ceramic part. Subsequent thermal processing decomposes the organic elements and induces sintering of the remaining elements. The thermal processes causes an inherent degree of shrinkage between the green and sintered components, typically 16 -18% [11], [12]. Additional finishing procedures388such as grinding, lapping and polishing can be used to improve the geometric tolerances and appearance of the component [13].

Additional finishing procedures388such as grinding, lapping and polishing can be used to improve the geometric tolerances and appearance of the component [13]. However, reliance on component specific templates and tooling place significant limitation on the flexibility and responsiveness of the manufacturing process. Moreover, tooling based production methods require volume production to be economically viable, whilst placing restrictive design constraints on the components that can be manufactured.Digital-driven manufacturing approaches such as additive manufacturing (AM) is an emerging field of manufacturing approaches that seek to mitigate the short-comings associated with traditional manufacturing approaches. Additive manufacturing produces components directly from digital data in a layer-wise manner, mitigating the need for component specific templates and tooling [14].

Additive manufacturing produces components directly from digital data in a layer-wise manner, mitigating the need for component specific templates and tooling [14]. Additive manufacturing was originally developed using polymers and waxes but have since been adapted to process materials such as engineering ceramics. Typically, this involves using conventional materials as a binding medium that is subsequently removed during thermal processing.Vat photopolymerization has received significant research attention owing to the fine features, low surface roughness and densities that can be achieved using this approach [15]. The process involves the homogeneous dispersion of ceramic material within the photocurable monomer that is selectively cured to fabricate the green ceramic part. Thermal processing of the part, decomposes the polymer and sinters the remaining material, yielding a monolithic ceramic component.

Thermal processing of the part, decomposes the polymer and sinters the remaining material, yielding a monolithic ceramic component. However, the additional material dispersed within the resin results in an alteration of the rheology and processing characteristics of the monomer. Consequently, the volume of ceramic that can be dispersed is limited, typically <41vol% [16] and shrinkages up to 30% can be experienced during thermal processing [17]–[19]. Moreover, the ceramic material modifies the interaction of the curing radiation and the monomer due to the different absorption characteristics of darker materials, thus restricting the range of materials that can be processed [20].Powder-based processes have also received significant research attention due to the potential applications for these types of processes including medical tissue scaffolds and process catalysts. Powder-based processes encompass Selective Laser Sintering/Melting (SLS/SLM) and Binder Jetting.

Powder-based processes encompass Selective Laser Sintering/Melting (SLS/SLM) and Binder Jetting. Whilst SLS/SLM has shown to be capable of directly processing ceramic material, issues with porosity, thermal shock and component integrity results in indirect processing methods being used predominantly [21]. Indirect fabrication techniques require the powder feedstock particles to be coated with a polymer coating to decrease sintering temperatures and reduce thermally induced stresses. Consequently, irregularities caused by the polymer coating reduces the flowability of the feedstock reducing density and minimum feature size [22], [23]. Alternatively, binder jetting doesn’t require modification of the powder feedstock as binder is selectively deposited onto the powder bed. The absence of heat during the fabrication of components mitigates thermally induced stresses whilst retaining flowability of the powder.

The absence of heat during the fabrication of components mitigates thermally induced stresses whilst retaining flowability of the powder. Despite this, all dry powder-based processes suffer from variation in the particle size distribution, preventing fully dense components from being fabricated and whilst components exhibit a porous and pitted appearance. Alternative, feedstock preparations have been investigated to address the issues with conventional feedstocks [24].Material extrusion is one most common AM processes owing to its relatively low cost and the numerous configurations available. The process involves the extrusion of material through a nozzle or orifice to form a continuous bead of material [25]. Material feedstocks can include389thermoplastic filaments that are highly loaded with ceramic materials [26], [27], ceramic slurries [28] or sol-gels [29]. The high viscosity of the feedstock materials facilitates higher loadings of ceramic material up to 75vol% [30].

The high viscosity of the feedstock materials facilitates higher loadings of ceramic material up to 75vol% [30]. Consequently, components experience smaller shrinkages compared to vat photopolymerization processes and components exhibit higher densities compared to powder-based processes. Moreover, the nature of the process offers greater multi-material compatibility compared to the aforementioned processes [31], [32]. Despite this, material extrusion processes are limited by manufacturing tolerances and poor surface finishes.Despite the relative merits of AM processes capable of processing ceramic materials, the inherent limitations of AM are compounded by the additional processing stages and associated shrinkage, resolution, density, material selection and compatibility. Hybrid manufacturing is an emerging class of digitally-driven manufacturing approaches in which multiple manufacturing approaches are amalgamated into a single, integrated system [33], [34].

Hybrid manufacturing is an emerging class of digitally-driven manufacturing approaches in which multiple manufacturing approaches are amalgamated into a single, integrated system [33], [34]. Hybrid manufacturing approaches that combine additive and subtractive manufacturing elements have been demonstrated in the production of customised large size ceramic moulds and templates using gypsum-based materials [35]. This paper presents the development of a digitally-driven hybrid manufacturing system for the direct production of components using engineering ceramics. Interleaving additive and subtractive elements into a single integrated system supported by auxiliary elements will facilitate the production of precision engineered components that exhibit low shrinkages and excellent multi-material compatibility.

Figure 1 shows the system that has been developed as part of this work.Sacrificial support printingMicro- machining spindleCeramic Paste depositionHeated, forced convection drying unitFigure 1 – Bespoke, digitally-driven manufacturing platform using a hybrid manufacturing approachMethodThe bespoke system has been developed for this application by interleaving a number of standalone subsystems into a single integrated manufacturing platform. The process is centred around an auger screw, high viscosity paste extruder (Eco-pen 300, Germany) using interchangeable luer lock nozzles. Subtractive machining is achieved using a precision micro-390machining spindle with automatic tool changer (Nakanishi, Japan) used in conjunction with fluted, CVD diamond machining tools. Auxiliary elements include a commercially available hot end (E3D, UK) with a 0.4mm nozzle and Bowden extruder (Bulldog Lite, UK) used to deposit PLA polymer.

Auxiliary elements include a commercially available hot end (E3D, UK) with a 0.4mm nozzle and Bowden extruder (Bulldog Lite, UK) used to deposit PLA polymer. Additionally, a heated, forced convection drying unit increases throughput of the process. The various elements have been mounted to a 3 axis CNC stage. The stage and the various elements of the process are actuated and controlled using commercially available control software that has been modified for this bespoke application (Newfangled Solutions, USA).The process was developed using 96wt% alumina with an average primary particle size of 2-12µm that had been formulated into a high viscosity, aqueous paste with a measured moisture content of between 18-22%, using a binding medium developed by Morgan Advanced Materials using a proprietary composition. Characterisation of the paste confirmed the pastes non-Newtonian properties and thixotropic response. This material was selected due to its broad application scope.

Characterisation of the paste confirmed the pastes non-Newtonian properties and thixotropic response. This material was selected due to its broad application scope. However, any ceramic material that can be formulated into a high viscosity paste with similar rheological characteristics can be processed. The rheology of the formulated material was investigated using an MCR 102 parallel plate rheometer (Anton Paar, Austria). Sheer values between 50-600s-1 were used. Figure 2 shows the results of the rheological tests which confirms the shear thinning rheological response of the precursor material, verifying its suitability for use in an extrusion type process.181614] s.

a P [ y t i s o c s i V121086420050100150200250 300 350 Shear Rate [1/s]400450500550600Figure 2 – Graph showing the results of the rheological tests on the pasteOptimisation of the deposition characteristics was undertaken using, 25 gauge (437µm) tapered nozzles with extrusion rates between 80 to 120µL/min and nozzle speeds between 380 and 470mm/min as these were found to produce the most consistent results. These parameters were identified as providing reasonable resolution and feature size with satisfactory production rates. Further optimisation of the processing parameters identified a nozzle speed of 410 mm/min with an extrusion rate of 104µL/min as being optimum for this process. The machining parameters were determined through the machining of components produced during these initial tests.

The machining parameters were determined through the machining of components produced during these initial tests. Ø4mm CVD diamond end mill was investigated using side and end milling operations with feedrates between 200-750mm/min with spindle speeds ranging from 6,000- 18,000 RPM. It was determined that a feedrate of 330mm/min and spindle speeds of 16,500 RPM achieved the best machining characteristics and lowest surface roughness. Figure 3 shows the process flow for the fabrication of a green state ceramic component.391Figure 3 - Process flow demonstrating the production of a green state ceramic componentThe fabrication of a green ceramic component involves the selective deposition of ceramic material with a controlled excess of around 2-5%, providing enough material to planarize the top surface and remove defects during post-process machining. A concentric infill pattern is used to prevent excess material within the perimeter regions of the component.

A concentric infill pattern is used to prevent excess material within the perimeter regions of the component. Deposition of material is proceeded by accelerated drying using the in-situ heated, forced convection drying unit, drying is undertaken until the component has a sufficiently low moisture content to facilitate machining. This is proceeded by a planarizing, machining operation to ensure flatness and consistency of the deposited layer. If required, this machining stage can also be used to improve the tolerance and fidelity of features, prior to the deposition of subsequent layers. Additionally, PLA can be deposited directly onto the machined ceramic which exhibits satisfactory adhesion to provide sacrificial support for overhanging features. During this work, investigation of enclosed sacrificial PLA supports was undertaken using structures of varying densities.

During this work, investigation of enclosed sacrificial PLA supports was undertaken using structures of varying densities. Additional post process machining operations can be undertaken using a variety of fluted machining tools to improve tolerances, remove defects and reduce the appearance of stair-stepping. Figure 4 shows the planarizing machining of a conformal geometry using the optimised machining parameters.Figure 4 - Interlayer machining of a conformal geometry using a Ø4mm CVD diamond end mill392Once fabricated, the components were thermally processed following a standard production thermal cycle in which temperatures exceed 1400°C. Testing of the manufactured substrates was undertaken to determine the density of the manufactured substrates, surface roughness and mechanical bend strength. Density and associated porosity was calculated using the Archimedean approach using a 0.1mg accuracy micro-balance (Sartorius, Germany) [36].

Density and associated porosity was calculated using the Archimedean approach using a 0.1mg accuracy micro-balance (Sartorius, Germany) [36]. Surface roughness measurements were obtained using 3-dimensional focus-variation measurements (InfiniteFocus, Alicona, Austria). Mechanical bend strength was determined by 3-point bend testing of samples measuring 5mm x 5mm x 50mm. The test samples were produced in a single batch to facilitate the direct comparison of results. The test parameters were derived from the standard tests used by Morgan Advanced Materials (Z005, Zwick/Roell, Germany).Results and DiscussionDuring the course of this work a bespoke, digitally-driven hybrid manufacturing platform has been developed for the production of high density ceramic components. A range of geometries have been fabricated including hemispheres, pyramids and conformal geometries.

A range of geometries have been fabricated including hemispheres, pyramids and conformal geometries. Additionally, components demonstrating internal cavities and voids have also been produced using PLA sacrificial support that was deposited directly onto the ceramic. Figure 5 shows some of the geometries that have been produced as part of this work.(a)(b)(c)Figure 5 – a: Green state Ø40mm conformal geometry having undergone post process machining operation. b: Conformal geometry post-thermal processing with a measured diameter of Ø33.2mm. c: a sintered pyramid, hollow hemisphere and tile that have been fabricated using the hybrid approach but have not been post processed using machining.A number of overhanging features and enclosed cavities have been demonstrated as part of this work. Figure 6a show the creation of internal cavities through the placement of393preformed support structures of varying densities.

Figure 6a show the creation of internal cavities through the placement of393preformed support structures of varying densities. Figure 6b shows the components the components once the support structure has been fully encapsulated within the ceramic components. Post process machining has been used to engrave the corresponding density into the ceramic component using the Ø1mm ball nose cutter. The ability to utilise a multitude of fluted and non-fluted machining tools improves the systems capability to produce high accuracy features. Consequently, finer features, smaller tolerances and greater fidelity with the digital data can be achieved compared to standalone AM systems. Figure 6c/d shows the components after undergoing thermal processing and the resultant internal cavity.(a)(b)(c)(d)Figure 6 - a: Varying infill densities of enclosed PLA support structures measuring 15mm × 15mm × 15mm.

b: PLA supports that have been fully encapsulated within the ceramic material and use of post process machining to engrave the corresponding densities of the support structure. c: The sintered samples, measured 12.3mm × 12.3mm × 12.3mm showed no signs of collapse, although delamination was evident on all samples. d: The internal cavity shows no signs of collapse or residue from the encapsulated support material.Analysis of manufactured components post thermal processing confirmed the production of monolithic ceramic parts that exhibited a measured shrinkage of ~18% between green and sintered parts. Visual inspection of the components revealed no visible residue attributable to the PLA support material. Samples with enclosed cavities containing PLA showed signs of delamination, possibly due to the expulsion of gases generated by the breakdown of the PLA support.

Samples with enclosed cavities containing PLA showed signs of delamination, possibly due to the expulsion of gases generated by the breakdown of the PLA support. Inspection of the cavity however, showed no signs of collapse with high fidelity to the cavity of the green state component. Additional testing of the manufactured components confirmed the production of full-density ceramic parts with part densities of ~99.8%. Surface roughness measurements of post machined parts exhibit a measured Ra of 1.11µm compared to measured Ra value of 8.09µm for non-machined parts. Mechanical bend strength was determined to be 218MPa. Minimum feature sizes were determined to be ±100µm. Figure 7 shows a SEM image of the interlayer region of a component, confirming the production of a monolithic ceramic component.

Minimum feature sizes were determined to be ±100µm. Figure 7 shows a SEM image of the interlayer region of a component, confirming the production of a monolithic ceramic component. The image confirms the complete sintering and adhesion between adjoining layers.394Figure 7 – SEM image of the interlayer region on a sintered partConclusionIn conclusion, this paper has presented the development of a digitally-driven, hybrid manufacturing platform for the production of precision engineered ceramic components of arbitrary complexity. The use of sacrificial PLA support structure has facilitated the production of overhanging geometries and enclosed cavities. The fabrication of components has been achieved using production methods comparable to conventional manufacturing approaches to yield monolithic components with densities in excess of 99.8% and 3-point bend strength of 218MPa and measured shrinkages of ~18%.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDIGITALLY-DRIVEN MICRO SURFACE PATTERNING BY HYBRID MANUFACTURINGMatthew A.A. Smith, Nicholas R Fry, Robert W Kay, Russell A HarrisFuture Manufacturing Processes Research Group, University of Leeds, Leeds, West Yorkshire, LS2 9JT UKAbstractAerosol Jet printing is a versatile direct-write method allowing selective deposition and alteration of surface chemistry on a variety of substrates, making it suitable for incorporation in a range of hybrid manufacturing processes. The digitally controlled nature of the presented hybrid manufacturing process enables rapid turnaround of designs, and improvements in flexibility and complexity compared to established methods.

The digitally controlled nature of the presented hybrid manufacturing process enables rapid turnaround of designs, and improvements in flexibility and complexity compared to established methods. The apparatus and instrumentation that has been created at the University of Leeds enables specific processing conditions that result in deposition of features with critical dimensions smaller than 20µm. In this study the analysis of the effect of process variables on deposition geometries is presented. The features were assessed by a combination of optical microscopy and white light interferometry. Using in- process machine vision, topographical compensation, and alignment capability the deposition of material into micropatterned features in poly(dimethylsiloxane) (PDMS) was demonstrated.

Using in- process machine vision, topographical compensation, and alignment capability the deposition of material into micropatterned features in poly(dimethylsiloxane) (PDMS) was demonstrated. High-value applications of this technology for surface functionalisation include electronics and bio-engineering.IntroductionAerosol Jet Printing is a promising new direct write (DW) technology that can be used to apply selective surface functionalisation. Aerosol Jet has been applied to a diverse range of applications, most notably in surface functionalisation and the prototyping of electronic circuitry [1]–[10]. In addition, the unique capabilities of Aerosol Jet have meant that it is integrated in several hybrid manufacturing processes chains [11]–[13], used in the production of embedded circuitry [14], and to integrated into bespoke machinery [15]–[17].

The ability to print microscale features on uneven surfaces [18], and the wide ranging choice of material [19]– [27] are compelling reasons to investigate the hybridisation of Aerosol Jet technology, as long as complementary processes can be identified and suitable automation procedures developed.In the Aerosol Jet technique, the functional material is transitioned from liquid ink into an aerosol state before being focussed and deposited onto a substrate surface (Figure 1). The material is atomised using either focussed ultrasonic energy, or pneumatic shearing of the fluid, depending on the viscosity of the liquid ink, and the required throughput of material for the application. The aerosol is entrained within a gas stream and transported to the deposition head by a stream of inert nitrogen gas [28]. The stream forms a boundary layer around the aerosol that prevents the majority of the particles in the aerosol stream from contacting the tubing during the transportation [29].

The stream forms a boundary layer around the aerosol that prevents the majority of the particles in the aerosol stream from contacting the tubing during the transportation [29]. An annular, co-axial sheath is introduced to the aerosol stream, which has the effect of collimating and accelerating the combined flow. The co-axial stream is further accelerated and focussed through nozzle directed at the substrate. Following the instructions contained in a digital script, the substrate is translated underneath the deposition head to produce traces of material. On/off patterning is achieved by interrupting the flow with a mechanical shutter.365Figure 1: Aerosol Jet Printing Overview. (1) A liquid sample is atomized by ultrasonic agitation. (2) N2 gas is pumped into the atomizer chamber. (3) The aerosol is transported to the deposition head (4) The aerosol is focused and accelerated by a further annular sheath of inert gas.

(2) N2 gas is pumped into the atomizer chamber. (3) The aerosol is transported to the deposition head (4) The aerosol is focused and accelerated by a further annular sheath of inert gas. (5) The resulting high velocity jet is deposited onto the substrate. (6) The stage is moved in up to 5 axes to produce a pattern. (7) On/off patterning is achieved by interrupting the flow with a mechanical shutterThe diameter and properties of the emerging stream (and therefore the geometrical properties of the deposit) are controlled by several adjustable process parameters. Mahajan et al. reported that in the case of printing silver nanoparticles, the critical factors in controlling the transportation and focussing mechanisms of the apparatus are the two gas flows and the movement speed of the substrate under the sample.

They contend that the total flow rate through a nozzle does not affect trace width, but that the ratio of sheath gas/atomiser gas, defined as the focussing ratio (Eqn 1) which is the critical factor. In addition, for each focussing ratio there is a range of optimum speeds at which a confluent line is drawn [30].𝑆𝑆ℎ𝑒𝑒𝑒𝑒𝑒𝑒ℎ 𝑔𝑔𝑒𝑒𝑔𝑔 𝑓𝑓𝑓𝑓𝑓𝑓𝑓𝑓 𝑟𝑟𝑒𝑒𝑒𝑒𝑒𝑒𝐹𝐹𝐹𝐹𝐹𝐹𝐹𝐹𝐹𝐹𝐹𝐹𝐹𝐹𝐹𝐹𝐹𝐹 𝑟𝑟𝑟𝑟𝑟𝑟𝐹𝐹𝐹𝐹 =𝐴𝐴𝑒𝑒𝑓𝑓𝐴𝐴𝐴𝐴𝑔𝑔𝑒𝑒𝑟𝑟 𝑔𝑔𝑒𝑒𝑔𝑔 𝑓𝑓𝑓𝑓𝑓𝑓𝑓𝑓 𝑟𝑟𝑒𝑒𝑒𝑒𝑒𝑒Printing materials must either be a liquid, or be dissolved, suspended or melted into liquid form. Their properties must then be tuned for Aerosol Jet printing, mainly through the addition of solvents, though more sophisticated material formulations can contain dispersants or other additives.

Their properties must then be tuned for Aerosol Jet printing, mainly through the addition of solvents, though more sophisticated material formulations can contain dispersants or other additives. Print materials are specifically designed for the process, considering: machine compatibility, stability through the printing process, and the final deposit properties.To produce a pattern, the substrate is manipulated under the printing head using two or more computer controlled axes. It is necessary to conduct a full analysis of the system, to achieve the functionality and repeatability required for integration into wider hybrid systems. Currently no publicly reported boundary conditions for the printing of geometrical features exist.In our work, we integrated the core Aerosol Jet technology into a bespoke programmable manufacturing process and used this to produce micro scale printed structures from the conductive polymer PEDOT:PSS.

This material was selected due to its range of high value applications such as touchscreens [31], solar cells [32], and as a non-toxic conductive material for electrical stimulation of cell cultures [33].366(1)A body of work is presented which:Identified trends in the printed line geometry according to process variables • Defined specific print parameters to print stable 20µm width at half maximum(WHM) features Investigated the machine and printing capabilities with a 20µm WHM lines including minimum line pitch, sharp angles, and minimum circle size. • Proved the application to complex surface patterning through example prints • Showcased the ability to combine patterned substrates through in processalignment.This core technology can be combined with multiple processes resulting in a multi- functional system.

Alignment capabilities enable integration with other manufacturing processes such as microstructures produced by photolithography.Studying the Effect of Printing VariablesThe integration of Aerosol Jet print engine into a bespoke multi axis stage resulted in a standalone manufacturing apparatus. The substrates were loaded onto the stage which is controlled through a G-Code input to Mach3 CNC software. The stage is actuated through lead screws driven by a geared stepper motor. It is capable of a minimum incremental movement of <10µm and speeds up to ~3mm/s (200mm/min) without loss of accuracy. The stage incorporates rotary alignment and an alignment camera capable of distinguishing features as small as 5µm. This enabled alignment to fiducial features on substrates if patterned substrates manufactured in other micromanufacturing methods (e.g. photolithography) were required in the application.

This enabled alignment to fiducial features on substrates if patterned substrates manufactured in other micromanufacturing methods (e.g. photolithography) were required in the application. The travel distance of the stage is 90mm x 90mm, enabling batch production of tests in a step and repeat fashion. The stage can be removed to transfer between pre- and post- processing steps.When printing, the ink was transitioned into an aerosol by ultrasonic atomization and deposited through a 100 µm diameter nozzle. The atomising current was kept constant at 0.65mA. The distance between the nozzle exit and the substrate was 3mm, and the substrate translated under the substrate at a speed of 2 mm/s.First, the geometry of the printed lines was assessed at different focussing ratios. Each print was conducted on a separate substrate to remove any chance of interference between prints.

Each print was conducted on a separate substrate to remove any chance of interference between prints. Before commencing each print, the aerosol jet apparatus was set to the specific sheath/atomiser gas flow rate and left to settle for 10 minutes to ensure a continuous stable flow from the nozzle. At this nozzle size the maximum recommended sheath gas is 50sccm.Table 1: Atomiser Gas flow rate (sccm) with respect to each Sheath Gas and Focusing Ratio combination.Focussing RatioSheath Gas (sccm)10 20 30 40 500.5 5 10 15 20 251 10 20 30 40 501.5 15 30 45 60 752 20 40 60 80 1002.5 25 50 75 100 1253 30 60 90 120 1503673.5 35 70 105 140 175Outside these regions no viable lines were printed. Focussing ratios below 0.5 resulted in machine issues such as nozzle clogging, focussing ratios above 3.5 resulting in very little material output, so that a confluent line was not deposited.

Focussing ratios below 0.5 resulted in machine issues such as nozzle clogging, focussing ratios above 3.5 resulting in very little material output, so that a confluent line was not deposited. To drive off solvents and sinter the nanoparticle line prior to analysis each sample was baked at 150°C for 10 minutes.Figure 2 shows a three-dimensional rendering of typical white light data, showing the curved cross-sectional profile of the deposits. The deposits do not have a well-defined sharp edge, and the side walls are not vertical. Therefore, the features were defined by their maximum height, and the width at half maximum.Figure 2: White light rendering a small section of aerosol jet deposit showing the cross sectional shape.The results of the geometrical analysis are presented in Figure 3. Figure 3a shows the effect of total flow rate on the maximum height of the deposit. At low focussing ratios the step height did not substantially increase with an increasing gas flow.

Figure 3a shows the effect of total flow rate on the maximum height of the deposit. At low focussing ratios the step height did not substantially increase with an increasing gas flow. When the focussing ratio is higher than 2 the maximum height increased as the total flow rate increased. Other than focussing ratio of 3.5, as the gas flow rate increased the maximum heights approach the same height, indicating a critical limit to the height of the printed features. Figure 3b shows that at all focussing ratios the WHM increased with an increasing gas flow. The most pronounced increase was seen at the lower focussing ratios, particularly those lower than 2.Generally, as the gas flow rates are increased, the material deposition rate onto the substrate surface is also increased, resulting in an enlargement of the deposited line. This tends to be in the form of an increase in height, until a critical point is reached, after which there is an increase in line width.

This tends to be in the form of an increase in height, until a critical point is reached, after which there is an increase in line width. Most notably, at a focussing ratio of 0.5 the height does not substantially increased, however large increased in WHM are observed as gas flow rate increased. Figure 3a indicates that the maximum deposit height for Aerosol Jetted PEDOT:PSS features is in the region of 0.5-0.8µm, as all the focussing ratios approach this region at high gas flow rates. Ultimately, this limit is as a result of the wetting properties of PEDOT:PSS to the substrate.368Figure 3 Effect of gas flow rate on (a) maximum height and (b) width at half maximum height.The effect of reducing atomiser gas whilst keeping the sheath gas constant can also be interpreted by comparing the points of each focussing ratio at a specific sheath gas value. For example, Figure 4 illustrates deposits at increasing focussing ratios at a constant sheath gas of 40sccm.

For example, Figure 4 illustrates deposits at increasing focussing ratios at a constant sheath gas of 40sccm. The reduction of atomiser gas at a constant sheath resulted in a reduced material flow rate from the atomising chamber, resulting in geometrically smaller deposits, both in terms of maximum height and width at half height.Sheath gas40 sccm40 sccm40 sccmFocussing ratio0.512Atomiser gas80 sccm40 sccm20 sccmFigure 4: Effect of decreasing atomiser gas at a constant sheath gas (S=40)There are limits and exceptions to these rules: in addition to the focussing ratio limits described earlier, several other sheath/atomiser gas combinations resulted in no printed features. This is due to the inability for the atomiser gas to push through enough material to form a confluent line. It is proposed that to define a printing width, first the atomiser gas should be set, so that a stable line is printed.

It is proposed that to define a printing width, first the atomiser gas should be set, so that a stable line is printed. The line geometry should then be altered within this range using the focussing ratio, and a knowledge of how the increased material deposition rate onto the substrate will affect line geometry.Validating Machine CapabilitiesSubsequent experiments investigated the capabilities of the bespoke automation system, working in tandem with the core aerosol jetting apparatus. The features that were assessed were the line pitch, sharp corners, and printing circles, as many complex shapes can be fundamentally made from these features. The aerosol was deposited through a 100 µm diameter nozzle. The36940 sccm313.3 scmmatomising current was kept constant at 0.65mA. The distance between the nozzle exit and the substrate was 3mm, and the substrate translated under the substrate at a speed of 2 mm/s.

The distance between the nozzle exit and the substrate was 3mm, and the substrate translated under the substrate at a speed of 2 mm/s. These experiments used a focussing ratio of 2, with the sheath gas set to 40 in order to print lines with a WHM in the region of 20µm. These extremely fine lines are desirable for many applications, particularly as the size of components in electronics and microfluidics becomes smaller. These experiments established boundary conditions for critical features when printing with a 20µm WHM line.Once a suitable line geometry was identified, and the process conditions established, the printed features were defined in G-code. Following the commands of the code, the substrate was manipulated in such a way to print discrete features. The use of a mechanical shutter to interrupt the flow enables the production of intricate, microscale patterns.Figure 5 shows the result of printing lines with decreasing pitch between samples.

The use of a mechanical shutter to interrupt the flow enables the production of intricate, microscale patterns.Figure 5 shows the result of printing lines with decreasing pitch between samples. The first observation was the starting bulge in all the prints, caused by the mechanical shutter opening before the stage was moved. This was due to the line by line execution method of G- code. Some advanced control codes enable the shutter to be opened independently of the stage movement, removing this artefact. Line integrity was confirmed by taking a cross sectional measurement. When the pitch of the line was smaller than two line widths (i.e. the space between the lines was smaller than the lines themselves), the regions of overspray noticeably interacted, and so the area between the lines was blurred. This has implications for applications where line integrity is key. For example, in fine pitch electronics this artefact may cause short circuits.

This has implications for applications where line integrity is key. For example, in fine pitch electronics this artefact may cause short circuits. For these applications is it recommended to print with a pitch greater than twice the line width.Figure 5: Pitch Experiments with a 20µm WHM line. (Left) White light interferometry measurements. (Right) Cross sectional profile measurement of white light data.Circles of decreasing radii were printed to determine the minimum radius to use in designs. In all these features it was noticed that the circle did not ‘complete’ and this was attributed to the backlash in the stage (~25µm in this system). Although backlash can be370designed out in many features, for circles it is unavoidable as both axes must always change direction to create it.

Although backlash can be370designed out in many features, for circles it is unavoidable as both axes must always change direction to create it. Some automation systems may include anti-backlash features such as using a ball screw, which would remove this artefact.Figure 6: Increasing diameter circles printed with a 20µm WHM line.When the same process conditions were used to print circles, the WHM increased slightly, due to the constant acceleration required to manipulate the substrate in the circular shape. Cross sectional profiles were taken to validate the geometry (Figure 7). The minimum diameter achievable by this system was roughly 100µm, however to print a truly circular feature the limit was 200µm. Below this diameter the effect of backlash was relatively large, and affected the circularity of the printed features.

Below this diameter the effect of backlash was relatively large, and affected the circularity of the printed features. As such a minimum recommended circle diameter of 200µm is recommended for a 20µm WHM line.Figure 7 Comparison of the profiles of the two smallest diameter circles. Circularity is affected by backlash in the 100µm diameter line.A series of sharp angles were printed to analyse the effect of a sharp corner on the depositions. Figure 8 shows the angles printed with 20µm width at half height line. In addition to the start and end bulges described previously, an accumulation of material was observed at the vertex of acute angles, due to the deceleration and acceleration profiles required to move the substrate through the sharp angle. As the angle increased and became obtuse, the stage does not need to change direction, so this effect was lessened.

As the angle increased and became obtuse, the stage does not need to change direction, so this effect was lessened. The manufacturing system successfully printed all corners, as sharp at 30°.371Figure 8: Increasing angle corners printed with a 20µm WHM lineSubstrate Alignment and Complex PatterningIntegration into hybrid manufacturing processes is reliant on the ability to precisely align substrates onto the machine. The in-process camera can align the substrate from a fiducial feature. The multi-axis stage incorporates one additional axis of rotation, meaning that the substrate can be loaded then aligned. The camera is capable of distinguishing features as small as 5µm, so substrates can be positioned with extremely high accuracy. The alignment capability was showcased by printing into 50µm channels, produced by casting from a photolithography mould, for extremely fine pitch microfluidic device.

The alignment capability was showcased by printing into 50µm channels, produced by casting from a photolithography mould, for extremely fine pitch microfluidic device. An optical microscopy image of the printed features is shown in Figure 9.Figure 9: Confluent PEDOT:PSS lines printed into 50µm wide channels in PDMS. Scale bar = 50µmAerosol Jet is suitable for integration in any process that is controlled by machine code. To change the pattern the code can be altered, without the need to update any templates. Therefore, even complex changes or new designs can be output quickly. A DXF to G-code converter was developed which enables complex designs to be incorporated into processing.

Therefore, even complex changes or new designs can be output quickly. A DXF to G-code converter was developed which enables complex designs to be incorporated into processing. Figure 10 showcases the ability to print intricate designs on a small scale by showing a University of Leeds logo printed within a 1.5mm square.Figure 10: Example of an intricate pattern produced by Aerosol Jet Printing (a) pattern viewed under an optical microscope with a match head for scale (b) white light interferometry scan showing fine details372ConclusionsIn summary, an automation system for Aerosol Jet printing has been characterised, and processing conditions to reliably print 20µm and smaller features from PEDOT:PSS were defined. At all focussing ratios, the material deposition rate was proportional to the total gas flow rate. The surface interactions of material and substrate affect the printed line geometry. In general, increasing the gas flow rate results in larger geometry.

The surface interactions of material and substrate affect the printed line geometry. In general, increasing the gas flow rate results in larger geometry. These increases manifested as either maximum height increase, or once a critical point is reached, they were due to an increase in line width. The approach to specifying a line width should be to first set an appropriate atomiser gas to achieve a stable deposit, followed by tuning the focussing ratio. It was recommended that for successful printing, the pitch of lines be no less than twice the line width, the diameter of a circle should be no less than 200µm, and any sharp corners should be an obtuse angle. The flexibility in terms of design was shown by printing the University of Leeds logo within a 1.5mm square. Aerosol Jet is an extremely versatile technology that can be applied to a range of scenarios.

Aerosol Jet is an extremely versatile technology that can be applied to a range of scenarios. Particularly, the opportunity for novel and multi-material printing, deposition onto non-planar substrates, and alignment capability enable integration within a wide range of hybrid processes that require selective surface alterations.AcknowledgementsThis work is funded by the UK Engineering and Physical Sciences Research Council (EPSRC) under grant EP/L02067X/2.References[1] A. J. Kell et al., “Versatile Molecular Silver Ink Platform for Printed Flexible Electronics,” ACS Appl. Mater. Interfaces, vol. 9, no. 20, pp. 17226–17237, May 2017.[2] B. Khorramdel, A. Torkkeli, and M. Mantysalo, “Electrical Contacts in SOI MEMSUsing Aerosol Jet Printing,” IEEE J. Electron Devices Soc., vol. 6, pp. 34–40, 2017.[3]J. Hester, E. Nguyen, J. Tice, and V.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDIMENSIONAL ACCURACY OF SMALL PARTS MANUFACTURED BY MICROSELECTIVE LASER MELTINGMichael Kniepkamp1, Jakob Fischer1, Eberhard Abele11 Institute of Production Management, Technology, and Machine Tools (PTW), Technische UniversitätDarmstadt, Darmstadt, GermanyAbstractWhile selective laser melting of metallic parts is already widely used in today’s industry, problems in this process still occur when using small parts with dimensions of less than 5 mm. Micro selective laser melting can fill gaps with layer sizes of less than 10 microns and powders with particle sizes smaller than 5 microns. In this paper the dimensional accuracy of parts with sub millimeter features using 316L steel powder is investigated.

In this paper the dimensional accuracy of parts with sub millimeter features using 316L steel powder is investigated. Test specimens with different features like slopes, overhangs and sharp radii were built applying different scan strategies. The parts were 3D scanned and compared to the CAD data to analyze their accuracy. Based on the results, optimized scan strategies for the different features were developed to increase the parts’ overall dimensional accuracy.1. IntroductionAdditive manufacturing (AM) is an emerging field in manufacturing technologies that has the common principle of building up solid parts directly from 3D CAD data by adding material layer by layer. Powder bed fusion based additive manufacturing processes use thermal energy to selectively fuse regions of a powder bed (ASTM F2792-12a 2012). Laser beam melting is a process where powder is applied in layers and then selectively melted using a laser beam to generate three-dimensional parts directly from CAD data.

Laser beam melting is a process where powder is applied in layers and then selectively melted using a laser beam to generate three-dimensional parts directly from CAD data. This study focuses on laser beam melting of metal powders, often termed Selective Laser Melting (SLM). Selective laser melting is typically involves layer thicknesses between 20 and 100 µm, using powders with particle sizes ranging from 20 to 45 µm (Gu et al 2012). The minimal layer thickness depends on the particle size distribution of the used powder. To increase the resolution and accuracy of selective laser melting, the process has been improved to realize the use of powders with smaller particle sizes enabling layer thicknesses of less than 20 µm (Streek et al 2014). Powders with mean particle diameters of less than 10 µm tend to agglomerate, which results in requiring new powder rake systems to process these materials.

Powders with mean particle diameters of less than 10 µm tend to agglomerate, which results in requiring new powder rake systems to process these materials. Since this process differs greatly from the already established SLM process in this study, the term Micro Selective Laser Melting (µSLM) is used. Dimensional accuracy in additive manufacturing describes the geometrical difference between the virtual designed CAD part and the physical part after the build process. Due to the layering manufacturing method, several effects have an influence on the dimensional accuracy of the parts. The machine accuracy describes the mechanical accuracy of the build system, like focused laser spot size, layer thickness and position accuracy of the scanner. The surface morphology, which can be characterized by several effects, also has an influence on the dimensional accuracy. These effects are highly dependent on the relative orientation of the part to the build direction (Strano et al 2013).

These effects are highly dependent on the relative orientation of the part to the build direction (Strano et al 2013). On top facing surfaces, the morphology is dominated by the stability of a single melt track and the hatch distance between two adjacent tracks (Yadroitsev and Smurov 2011, Pupo et al 2014). As the top layer is always accessible by the laser, re-melting strategies can be applied to reduce the surface roughness during the build process (Yasa and Kruth 2011). The morphology of side facing surfaces, which are parallel to the build direction, is dominated by partial melted powder particles. Side facing surfaces are surrounded by loose powder particles during the build process. These particles get drawn into the melt pool but due to insufficient energy on the melt pool edge only partial melting occurs. This effect can be influenced by the process parameters laser energy and scan speed (Mumtaz and Hopkinson 2009).

This effect can be influenced by the process parameters laser energy and scan speed (Mumtaz and Hopkinson 2009). The third effect on dimensional accuracy is the distortion of the parts caused by anisotropy and residual stresses. These occur due to high temperature gradients as a result of the locally concentrated energy distortion which may occur when removing the part from the substrate on which it is produced (Kruth et al 2010). In selective laser melting different exposure steps are being used. A common strategy is to use two exposure 11530steps; one for the contour of the part, and one for the core (Meiners 1999). The contour scan vector and its width determines the build accuracy. An beam offset between the parts contour and the contour vector is used to compensate the width of the contour track (Abele et al 2015). Benchmark parts are commonly used to measure the machine and process accuracy.

Benchmark parts are commonly used to measure the machine and process accuracy. Benchmark parts usually have different features, such as slopes, thin walls, overhangs and sharp corners (Byun and Lee 2003, Mahesh et al 2004, Kruth et al 2005). These parts have been designed so that they do not need a large amount of material in order to be built, to not take too long to build and to be easily and quickly measurable by coordinate measuring machines. Due to their size, up to this date benchmarking parts have not been possible to use for the µSLM process. In this study, the dimensional accuracy of the µSLM process is measured using a new test specimen with a maximal diameter of 5 mm based on the typical size of real parts for µSLM.2. Research methodologyA commercially available µSLM system EOSINT µ60 from the 3D-Microprint GmbH (Germany) was used to conduct the experiments. This system uses a 50 W single mode fiber laser with a wavelength of 1060 nm focused to a spot size of 30 µm.

This system uses a 50 W single mode fiber laser with a wavelength of 1060 nm focused to a spot size of 30 µm. The laser can be operated in a pulsed or continuous wave mode. In this study, the laser was operated only in the continuous mode, as the pulsed mode leads to discontinuous melt tracks resulting in a high surface roughness. The build platform of the system has a diameter of 60 mm and is moved by piezo actuators with an accuracy of less than 1 µm. To apply layer thicknesses of less than 10 µm powders with respectively small particle, diameters have to be used.Table 1 Reference exposure parametersExposure StepOperation modeLaser powerScan speedBeam offset1. Corecontinuous wave40 W1000 mm/s58 µm2. Contourcontinuous wave40 W500 mm/s26 µmThese powders tend to agglomerate. This makes it impossible to coat the powders using gravitational forces only.

Contourcontinuous wave40 W500 mm/s26 µmThese powders tend to agglomerate. This makes it impossible to coat the powders using gravitational forces only. Thus powder application is achieved by pressing and wiping the powder onto the build platform using external force. The powder used is a 316l gas atomized stainless steel powder with D50 = 3.62 µm and D90 = 5.69 µm. Previous studies using the µSLM system optimized the build strategy to high relative densities of more than 99 % (Fischer et al 2014) and low surface roughness. The result of this is a two-step exposure strategy for the part core and the part contour (Table 1).

The result of this is a two-step exposure strategy for the part core and the part contour (Table 1). The parameter setup has a mean roughness Ra on vertical surfaces of 1.69 µm (Abele and Kniepkamp 2015) and less than 1 µm on top surfaces.Figure 1 Test specimen to analyze dimensional accuracy (Units: mm)15312To analyze the dimensional accuracy of the µSLM machine, a test specimen based on a typical use case for this process has been used (Figure 1). The part has a tube shaped base plate with an outer diameter of 5 mm and a height of 1 mm. 4 angled pins with heights of 0.5 and 1 mm are on the top. The baseplate is used to analyze larger volumes, outer (5 mm) and inner (1.5 mm) diameters and 45° slopes. The pins are used to measure the accuracy in Z-direction, smaller volumes, small radii (0.1 mm, 0.2 mm and 0.03 mm) and 75° overhangs.

The pins are used to measure the accuracy in Z-direction, smaller volumes, small radii (0.1 mm, 0.2 mm and 0.03 mm) and 75° overhangs. To analyze the dimensional accuracy, the specimens were separated from the build platform and three dimensionally scanned using an Alicona InfiniteFocus optical measurement system with an objective magnification factor of 5. Focus variation with a resolution of 410 nm and a rotational unit is used to get a full 360° scan of the test specimens. The software GOM Inspect is used to compare the scanned 3D model with the original CAD designed STEP file.

The software GOM Inspect is used to compare the scanned 3D model with the original CAD designed STEP file. Within GOM Inspect the overall surface deviation and discrete points along critical features of the test specimen are analyzed (Figure 2).Figure 2 Critical features on the test specimen surfaceTo determine the reproducibility of the µSLM system and the dimensional accuracy of the initial exposure strategy (Table 1), 10 specimens were built in two operations, both with an even distribution of the parts across the build platform. To further increase the dimensional accuracy of the initial exposure strategy, further experiments with different strategies were carried out. To increase the accuracy of the pins and the overhangs, the laser power has been reduced to decrease the energy and achieve smaller melt pools.

To increase the accuracy of the pins and the overhangs, the laser power has been reduced to decrease the energy and achieve smaller melt pools. The energy required to generate parts in selective laser melting can be described with the volume energy (𝐸𝑉) (Meiners 1999), with the laser power (P) the scan speed (v) the hatch distance (h) and the layer thickness (d).𝐸𝑉 =𝑃 𝑣 ℎ 𝑑The laser power was reduced in 5 W steps to the lower end of the stable line energy of the process window for this material. The hatch distance and the layer thickness were kept constant at 33 µm and 7 µm using a scan speed of 1000 mm/s. The resulting volume energies for can be seen in Table 2.Table 2 Volume energies at reduced laser powerLaser power [W]4035302520Volume energy [J/mm³] 173.16151.52129.87108.2386.581532(1)3With a decrease in volume energy, an increase of porosity within the parts is to be expected.

To analyze the porosity cubes with lateral lengths of 5 mm have been built using the lowest and the highest volume energies of 173.16 and 86.58 J/mm³. The density was determined using the Archimedes weighting. A downskin build strategy was used to increase the accuracy of the overhanging regions. Separate exposure strategies for the overhanging structures and the rest of the part were used. As the CAM software used for µSLM does not currently support separate scan strategies for different areas of a part, two parts were generated; one for the overhanging structure, and one for the rest of the part (Figure 3). The width of the downskin structure was set to 0.1 mm and the top layer was excluded from the overhang part. To increase the accuracy of the overhanging structure, the volume energy was reduced to 86.58 J/mm³ and the distance to the part contour was compensated by 19 µm which is the mean deviation of the reference parameter set.

Each specimen was built 5 times and evenly distributed on the build platform.Figure 3 Specimen for overhanging structures. Overhanging structure (red)3. ResultsThe results of the dimensional accuracy of the reference parameter setup are summarized in Figure 4. The mean deviation of the inner and outer diameters of the baseplate is within a range of ± 6 microns. In comparison to these results, the deviations of the two chamfers were found to be high with a mean deviation of -17.8 microns and peaks of more than 40 µm. The two larger radii (0.2 mm and 0.1 mm) of the pins and the sloped surface had low deviations the overhanging surface and the small radius show high deviations of 19.6 and -26.4 microns. The highest deviations from the designed CAD part can be seen on the top surface of the pins. The pin top contour had peak deviations of more than 60 microns in both directions and the mean deviation of the center surface is 13.6 microns higher than designed.

The pin top contour had peak deviations of more than 60 microns in both directions and the mean deviation of the center surface is 13.6 microns higher than designed. The standard deviation of the analyzed features occurred in the layer thickness of 7 microns. However, the pin contour and the overhanging surface had a standard deviation of more than 20 microns.FeatureMean deviation [µm]Std. dev. [µm]Inner diameter5.43.6Outer diameter4.15.7Pin contour8.220.5Pin center13.66.830 µm radius26.46.6Chamfer17.86.7Overhang19.620.6Figure 4 Dimensional accuracy of the reference parameter setup. Mean deviation of the features (left) example of overall surface deviation (right).15334The results of the volume energy variation are summarized in Figure 5. Lower volume energies of the parts’ core exposure have no significant influence on the outer and inner diameters of the base plate or the 30 microns radius and the overhanging surface.

Lower volume energies of the parts’ core exposure have no significant influence on the outer and inner diameters of the base plate or the 30 microns radius and the overhanging surface. The dimensional accuracy of the chamfers and the top center surface of the pins is increased significantly.20 W 25 W 30 W 35 W Reference (40 W)3020]m µ [ n o i t a i v e d n a e M100102030Inner diameter Outer diameterPin conturPin center30 µm radiusChamferOverhangFigure 5 deviations with decreasing volume energiesStd. dev.Using a volume energy of 86.58 J/mm³, the mean deviation of the chamfer can be reduced from -17.8 down to -5.1 microns and on the pin center from 13.6 to 6.9 microns. The mean deviation of the pin contour shifts from -8.2 to 11.1 microns. The standard deviation of the pin contour is between 14 and 20 microns across all laser powers.

The mean deviation of the pin contour shifts from -8.2 to 11.1 microns. The standard deviation of the pin contour is between 14 and 20 microns across all laser powers. The density of the high volume energy (173.16 J/mm³) is at 7.9 g/cm³ and 7.7 g/cm³ for the low volume energy (86.58 J/mm³). The relative density decreased to 97 % compared to the reference parameter set (100 %). The deviation of the overhang surface using the downskin build strategy can be seen in Figure 6. The mean deviation in the overhanging structure is reduced from 19.5 microns to 5.5 microns. While the deviation on the flat overhang structure reduced the deviation, the partly overhanging curved surfaces adjacent to it, there was a deviation increase of up to -43 microns.3025]m µ [20n o i t a i v e D n a e1510M50ReferenceCompensated OverhangFigure 6 Overhang deviation of the downskin build strategy4.

DiscussionTo increase the resolution of the selective laser melting process, new machines with higher precision are needed. In theory, reducing the layer thickness and focus diameter should enable generation of smaller parts with a high accuracy. The reference parameter set for the µSLM system was previously optimized towards high densities and low surface roughness. In this study, firstly dimensional accuracy of small parts with dimensions of less that 5mm was investigated. Using the reference parameter setup some areas of the test specimen, such as the base plate, can be built with an accuracy of ± 10 microns. The mean surface roughness for the reference parameter set is lower than 2 microns on the top and side facing sides. In theory, the best possible dimensional build accuracy is within the surface roughness. In reality, other effects also have an impact on the accuracy of additive manufacturing processes.

In theory, the best possible dimensional build accuracy is within the surface roughness. In reality, other effects also have an impact on the accuracy of additive manufacturing processes. Melt pool dynamics, surface orientation, the coating process and the parts geometry also have an influence on the overall dimensional accuracy. The smallest feature on the test specimen, with a radius of 30 microns, deviate greatly to the designed CAD part. The width of the reference parameters contour 51534Figure 7 SEM images of the top pin surface with 40 W (left) and 20 W (right) core laser power.scan track is 52 microns. To compensate the scan track width, a beam offset of 26 microns is used. In theory, the smallest possible radius using this compensation is also 26 microns which is lower than the designed 30 microns of the test specimen. The reason for the negative deviation of the small radius can be explained with the SEM images of the pins top surface (Figure 7).

The reason for the negative deviation of the small radius can be explained with the SEM images of the pins top surface (Figure 7). The used CAM software sets the starting and end points of the contour vector on the small radius of every layer [1]. The disconnected melt pool on the edge creates defects and, as a result, the part becomes smaller than the designed size. The deviations on the top pin contours can also been seen in Figure 7. The SEM images reveal two defects in the contour area. The first one is an insufficient connection between the contour and the core vectors [2]. This results in gaps between the two exposure steps. This can be compensated by decreasing the offset between the core and the contour vectors. The second observed phenomenon is an irregular track width of the contour vector [3]. While it is constant on the straight areas of the contour, it is significantly thinner in the curved areas.

The second observed phenomenon is an irregular track width of the contour vector [3]. While it is constant on the straight areas of the contour, it is significantly thinner in the curved areas. Due to the use of the STL file format, the curved areas of a part are triangulated to smaller straight lines instead of a continuous curve. The laser scanner has to rapidly change direction when scanning those smaller lines. This leads to smaller melt pools and results in a thinner track width with a decreasing curve radius. The variation of the part’s core volume energy results in an increase of dimensional accuracy on the pin’s top surface and the chamfers. The decrease of the laser power, while keeping the hatch distance, results in thinner track width and gaps between the cores scan vectors (Figure 7) [4]. Less material is molten, which reduces the positive deviation from the designed part on smaller cross sections.

Less material is molten, which reduces the positive deviation from the designed part on smaller cross sections. A disadvantage of this strategy is the reduced density due to pores in the part and a decrease of the top surface roughness. In theory, heat dissipation on overhanging structures is lower due to the lower thermal conductivity of the loose powder below the overhanging structure. This effect can be compensated by reducing the volume energy (Mertens et al 2014). The deviation in the overhanging parts of the pins cannot be reduced with lower volume energies of the core exposure and constant line energies of the part’s contour. With an overhanging angle of 75° and a layer thickness of 7 microns, the theoretical horizontal down facing area in each layer is 1.87 microns. With a track width of 52 microns, the contour exposure step dominates the core exposure in the overhanging area.

With a track width of 52 microns, the contour exposure step dominates the core exposure in the overhanging area. Reducing the laser power of the contour vector could increase the dimensional accuracy in the overhanging structure with the drawback of a higher surface roughness in on the whole part. Using the down-skin exposure strategy the deviation in the overhanging structure could significantly be reduced by adjusting the contour offset. The15356disadvantage of this method is that a different offset is needed for each angle. Another challenge is to generate a smooth transition between the overhanging areas and the rest of the part.5. Summary and outlookThe present work gives a first impression of the dimensional accuracy of small metallic parts built using micro selective laser melting with layer thicknesses of 7 microns. Test specimens with a maximum dimension of 5 mm have been built using different exposure strategies.

Test specimens with a maximum dimension of 5 mm have been built using different exposure strategies. The specimens were 3D scanned using optical focus variation and compared to the designed geometry. From the results the following conclusions can be drawn: Using a reference exposure strategy optimized for density and surface quality, mean deviations from the designed geometry ranging from -26.4 to 20 microns with peaks of ± 60 microns can be seen. The built accuracy of larger cross sections is higher than for smaller ones. The highest deviationscan be seen on overhanging structures and on the top surface of small cross sections.  The accuracy of small cross sections can be increased by decreasing the laser power.  Using different exposure strategies for the overhanging structures and the rest of the part can beused to increase the accuracy of overhanging areas.To achieve the best build accuracies in all areas of a part, a uniform exposure strategy should not be used.

Future research should analyze in detail which geometrical features require special exposure strategies in order to further increase the dimensional accuracy of µSLM parts. Furthermore, an automated process needs to be developed, in order to assign different exposure strategies to different areas within one part.AcknowledgmentsThis work was partially funded by the Federal Ministry of Economic Affairs and Energy through theAiF GmbH (Grant No. KF2012461WO3). The Authors would like to thank for this support.ReferencesAbele E and Kniepkamp M 2015 Analysis and optimisation of vertical surface roughness in micro selectivelaser melting Surf. Topogr.: Metrol. Prop.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDIMENSIONAL ACCURACY OF TITANIUM DIRECT METAL LASER SINTERED PARTSW. F. Mitchell*†, D. C. Lang†, T. A. Merdes†, E. W. Reutzel*†, and G. S. Welsh§Center for Innovative Material Processing through Direct Digital Deposition (CIMP-3D) † Applied Research Laboratory at the Pennsylvania State University, University Park, PA 16801 § Naval Air Systems Command, Patuxent River, MD 20670AbstractTo address concerns regarding quality of production parts created using the Additive Manufacturing (AM) process, a study was conducted to quantify the dimensional accuracy of said parts. Fourteen AM builds were manufactured in Ti-6Al-4V material across two EOS DMLS machines (EOSINT M 280 and EOS M 290).

Fourteen AM builds were manufactured in Ti-6Al-4V material across two EOS DMLS machines (EOSINT M 280 and EOS M 290). In addition to studying the impact of machine-to- machine variability, other factors potentially impacting final dimensional accuracy were studied, including: powder state (virgin or reused); post-processing steps (heat treatment and part removal from substrate); location of part on substrate; and nominal part size. The results of the dimensional analysis showed that the individual machine itself was the dominant factor impacting dimensional accuracy. Also, a non-linear relationship between dimensional accuracy and nominal part size was identified, which would require a more complex machine calibration technique to correct.IntroductionManufacturing metal parts using the Additive Manufacturing (AM) process is of greatinterest to the aerospace industry [1].

The AM process has many potential advantages compared to conventional manufacturing processes including additional design freedom, creation of internal geometry, complex light-weighting of designs, consolidation of assemblies, and a reduced logistical supply chain [1]-[2]. In addition, certain titanium alloys are readily processed using AM, and the resultant parts offer potential to be incorporated into complex aircraft designs [1]. In order to begin using AM to produce production parts in these applications, an extensive effort must be undertaken to qualify and certify the process and final parts [2]. Dimensional accuracy must be a part of qualification and certification to ensure the parts being created on an AM machine match the final design intent. Also, to realize the economic benefits of AM [2], post processing must be minimized.

Also, to realize the economic benefits of AM [2], post processing must be minimized. Though many current applications require the use of conventional subtractive techniques to achieve the final part tolerances, particularly surface roughness [3], as more cost efficient techniques (e.g. batch surface finishing operations such as electrochemical polishing, Micro Machining Process, and Isotropic Superfinishing) become established to meet surface roughness limits, dimensional accuracy will be the next limiting factor in achieving final part tolerances.

A combination of new surface finishing techniques and tighter, quantifiable dimensional control promises to reduce or possibly eliminate post- processing steps such as conventional machining, thereby significantly reducing final part cost.Towards ultimately reducing final part cost and establishing part certification, a primarygoal of our ongoing research is to investigate the dimensional accuracy that can be realized on2029commercially available AM machines. This paper covers the present state of that investigation. The research investigates Ti-6Al-4V parts that were produced on two EOS Direct Metal Laser Sintered (DMLS) machines located at different facilities. The final dimensional accuracy analysis includes an assessment of the overall dimensional accuracy realized; an investigation into variables possibly influencing the dimensional accuracy; and proposals for increasing the process accuracy.Experimental SetupFourteen AM builds were produced for analysis.

All builds used identical layouts, parameters, and procedures. Every effort was made to have consistency between each build and eliminate unintended sources of error. The builds were distributed across two machines and also encompassed virgin (previously unused) and reused titanium powder.The build layout contained production components, mechanical test specimens, and so- called “witness coupons” (see Figure 1). The witness coupons were designed to assess various build characteristics, including dimensional accuracy. Both the mechanical test specimens and witness coupons were measured upon completion of each key manufacturing process which included the AM build, hot isostatic pressing (HIP), and part removal via submerged wire electrical discharge machining (EDM). In accordance with ASTM F2924-14 [4], the substrate with parts still attached was HIPed in an argon environment at 900 °C and 102 MPa for 4 hrs.

In accordance with ASTM F2924-14 [4], the substrate with parts still attached was HIPed in an argon environment at 900 °C and 102 MPa for 4 hrs. A submerged wire EDM was then used to separate the parts from the substrate, cutting as close to the substrate as possible. In this application, the AM parts are ready for final finishing and/or machining upon completion of the three manufacturing processes listed above.Figure 1: Part layout on AM build (left). Both mechanical test specimens and witness coupons measured for dimensional analysis (right).The two AM machines used in this research included an EOSINT M 280 and an EOS M 290. Both machines are based on Direct Metal Laser Sintering (DMLS) technology, the EOS M 290 being a newer version of the EOSINT M 280 with some enhancements [5]. The typical achievable part accuracy in Ti-6Al-4V material is stated as equivalent between the M 280 and M 290 machines [6], and they were treated as equal in this study.

The typical achievable part accuracy in Ti-6Al-4V material is stated as equivalent between the M 280 and M 290 machines [6], and they were treated as equal in this study. Both machines were calibrated by EOS technicians for laser power, focal distance, and galvanometer accuracy during installation.2030Six AM builds were completed on the EOSINT M 280 machine and eight AM builds were completed on the EOS M 290 machine.Ti-6Al-4V metal powder was used for all AM builds. The raw material was sourceddirectly from the machine manufacturer [6]. Half of the builds from each machine (three for the M 280 and four for the M 290) used powder which had never undergone AM processing, otherwise known as virgin powder. After successful completion of each virgin powder build, the remaining unsintered powder was reclaimed and sieved according to the manufacturer’s recommendations.

After successful completion of each virgin powder build, the remaining unsintered powder was reclaimed and sieved according to the manufacturer’s recommendations. The reclaimed powder, now designated as “reused,” was then used to complete another round of AM builds on both the M 280 and M 290 machines. The effect of powder state, virgin vs. reused, on dimensional accuracy is analyzed in the results section.For dimensional accuracy assessment, the aforementioned mechanical test specimens and witness coupons were measured with calibrated calipers after each key manufacturing process, as described above. Exceptions where measurements were unable to be collected included M 280 builds 5 and 6, post part removal and all M 290 builds, post build. The same technician measured all parts to eliminate user variability from influencing the measurement results. The measurements were recorded into a database for storage and future analysis.

The same technician measured all parts to eliminate user variability from influencing the measurement results. The measurements were recorded into a database for storage and future analysis. The machine, powder state, and manufacturing step were all analyzed for their influence on dimensional accuracy. Additionally, the final part measurements (after part removal) were analyzed for correlations to part location on the substrate and nominal part size.ResultsOnce all fourteen AM builds were completed and part measurements collected, the partmeasurement data was normalized for dimensional analysis.

reused), and post-processing step. The dimensional accuracy data was also analyzed for trends with respect to part location on the substrate and nominal part size. The two regression analyses were used to characterize changes in dimensional accuracy caused by build quality variations or machine calibration settings (e.g., laser beam offset and X-direction or Y-direction scaling errors).The build-to-build variation was analyzed first. For comparison, the data was restricted to measurements taken after part removal. Also note-worthy, builds 1-3 on the M 280 machine and builds 1-4 on the M 290 machine were completed using virgin powder. The remaining builds were completed with reused powder.Utilizing a box and whisker plot, Figure 3 visually summarizes the dimensional accuracy variability from build-to-build on both the M 280 and M 290 machines.

In box and whisker plots, the upper and lower extremes of the boxes represent a given population’s interquartile range (25th percentile to the 75th percentile); the whiskers extending from the interquartile boxes represent the upper and lower quartiles; and the asterisks represent outliers in the data. Within the interquartile box, the population’s median is shown as a horizontal line and the population’s mean is shown as a filled diamond.

Within the interquartile box, the population’s median is shown as a horizontal line and the population’s mean is shown as a filled diamond. The means are connected from build-to-build within each machine to highlight dimensional accuracy changes with respect to build number.0.2Machine M 280 M 290)m m( y c a r u c c A0.10.0l a n o i s n e m Di0.10.2Build #1B uild #2Build #3B uild #4B uild #5B uild #6B uild #7B uild #8Figure 3: Box and whisker plot comparing dimensional accuracy across multiple AM builds and two AM machines.An analysis of variance (ANOVA) was performed on the build-to-build data to quantify any statistical difference between each build. Results from a pairwise comparison using Tukey’s method at 95% confidence level showed that there was no statistical difference between the means of all builds, when restricted to each machine separately.

Both the M 280 and M 2902032machines produced parts on a build-to-build basis that were dimensionally equivalent at 95% confidence. Therefore, the data suggests there was no statistically significant drift of the dimensionally accuracy from build-to-build on either machine. [7]As is suggested in the build-to-build data (Figure 3) and build-to-build pairwise comparison results, the builds completed on the M 280 machine produced dimensional accuracy results that differed from builds completed on the M 290 machine. To statistically assess this study’s second potential factor, machine-to-machine variability, the dimensional accuracy data were segregated into separate populations for each machine. A comparison of each machine’s dimensional accuracy distribution can be seen in the histogram below (Figure 4). The histogram shows that the two machines produced dimensional accuracy populations with similar variance (σ2), but different means [7].

The histogram shows that the two machines produced dimensional accuracy populations with similar variance (σ2), but different means [7]. The M 280 and M 290 data’s standard deviation was σ = 0.061 mm and σ = 0.062 mm respectively. However, the data’s means were μ = 0.038 mm and μ = -0.050 mm respectively, a total difference of 0.088 mm.21enihcaM 082 M 092 M018y t i s n e D642081.0-21.0-60.0- D00.060.0)mm( ycaruccA lanoisnemi21.0081.Figure 4: Histogram of dimensional accuracy differences between two AM machines.Using the two-sample t-test for equal means, the M 280 and M 290 dimensional accuracymeans were statistically compared. At a 95% confidence level, the t-test rejected the null hypothesis signifying that the each machine’s mean dimensional accuracy was statistically different. From the two-sample t-test, the test statistic (t-value) was computed to be 15.30 and the calculated probability (p-value) was computed to be less than 0.000.

From the two-sample t-test, the test statistic (t-value) was computed to be 15.30 and the calculated probability (p-value) was computed to be less than 0.000. Though standard deviations for both machines were nearly identical, both calculated t- and p-values exceed the limits to accept the null hypothesis at 95% confidence (less than 1.96 for t-value and greater than 0.05 for p-value), validating that the means between each machine are different with statistical significance. [7]Powder state was the third factor analyzed for its effect on dimensional accuracy. Anequal number of AM builds were completed on the M 290 machine for each case: four utilizing2033virgin powder and four utilizing reused powder. To assess the effect of powder state on the dimensional accuracy while limiting extraneous influences, the data was restricted to include measurements captured after part removal, from M 290 AM builds only.

The machine type and manufacturing processes are excluded from the analysis, thereby eliminating their possible influence on the results.The virgin and reused powder dimensional accuracy results can be seen in Figure 5. Thestandard deviation of each population was σ = 0.062 mm and σ = 0.059 mm for virgin and reused powder respectively. The population means were μ = -0.050 mm and μ = -0.060 mm respectively. Again using the two-sample t-test for equal means, the virgin powder and reused powder populations were statistically compared. The resulting test statistic and calculated probability were computed to be t-value = 1.77 and p-value = 0.077. Both values fall within the limits to accept the null hypothesis, H0, at a 95% confidence level.

Both values fall within the limits to accept the null hypothesis, H0, at a 95% confidence level. Therefore, the two-sample t- test for equal means indicates that there is no statistical difference in dimensional accuracy between the virgin powder and reused powder population means.0.10)m m0.05( y c a r u c c A0.000.05l a n o i s n e m Di0.100.150.200.25VirginOnce UsedPowder StateFigure 5: Box and whisker plot comparing dimensional accuracy between virgin and reused powder. Statistical testing indicated no difference between the two population means.The fourth and final statistical comparison assessed the effect of each manufacturingprocess on dimensional accuracy. Understanding the effect, if any, caused by HIP or part removal on dimensional accuracy would allow the initial AM part design to be compensated accordingly. For this analysis, the data was filtered to include measurements taken on M 280 AM builds using virgin powder only.

For this analysis, the data was filtered to include measurements taken on M 280 AM builds using virgin powder only. Filtering in this manner limits any extraneous effects from machine-to-machine and powder state variability.The dimensional accuracy results can be seen in Figure 6 below. The three box andwhisker groups represent the dimensional accuracy measurements taken after each manufacturing process. An ANOVA was completed on the three data groupings, the results of2034which indicated no statistical difference in the variances and means when comparing “Post Build” to “Post HIP” and “Post HIP” to “Post Part Removal”. The t-test calculated the t-values to be 1.51 and 1.49 respectively while the p-values were calculated to be 0.285 and 0.296 respectively. However, the ANOVA results did indicate a statistical difference of the variances and means between the “Post Build” and “Post Part Removal” groups, with the t-test calculating a t-value of 3.03 and a p-value of 0.007.

This appears to indicate that the part size is systematically reduced once processed through HIP and part removal via wire EDM. Since the total difference between the “Post Build” and “Post Part Removal” means was 0.032 mm while the precision uncertainty of the measurement device was calculated to be 0.010 mm, collecting more data is desired before drawing definitive conclusions.0.50.4)m m(0.3y c a r u c c A0.2l a n o i s n e m Di0.10.00.10.2Post BuildPost HIPPost Part RemovalStateFigure 6: Box and whisker plot comparing the dimensional accuracy across three key manufacturing processes.With all categorical factors analyzed, the study was expanded to characterize dimensionalaccuracy variation as a function of two continuous variables: part location on substrate and nominal part size.

For both analyses, only measurements taken after part removal were used.First, a linear regression was performed on the data with dimensional accuracy as thedependent variable and linear distance from the substrate center to part location as the independent variable (see Figure 7). A scatter plot of the data and regression analysis results can be seen in Figure 8. Prior to analysis, the dimensional accuracy was hypothesized to change as the part under measurement increased in distance from the substrate center due to expected variations in laser spatial energy distribution across the build area. The largest laser variation should occur in the corners of the building area, therefore those areas have the highest probability of large dimensional deviations.2035Figure 7: Illustration of various distances from the center of substrate.

Dimensional accuracy measurements were analyzed according to the distance from the center of the substrate to the centroid of the part under measurement.Figure 8: Dimensional accuracy plotted as a function of distance from substrate center to part measurement location.2036The fitted coefficients for the linear equation of the form (cid:1877) (cid:3404) (cid:1865)(cid:1876) (cid:3397) (cid:1854) are shown in Table1. Table 1 also summarizes the goodness of fit value (coefficient of determination or R2 value) and results from the F-test (F-value and p-value). The F-test compares the linear regression model against the regression model one degree lower, in this case a constant model. The F-test outputs a test statistic, the F-value. If the F-value exceeds a threshold value established by the number of degrees of freedom and confidence level, the linear regression model is said to fit the data better than the constant model.

If the F-value exceeds a threshold value established by the number of degrees of freedom and confidence level, the linear regression model is said to fit the data better than the constant model. The p-value is an indicator associated with the F-value and when it is below the desired significance level (in this case, α = 0.05), the regression model is considered to be a better fit to the data, with statistical significance, than a regression model one degree lower. dimensional accuracy and part location. The regression analysis confirms these visual observations. For both machines, coefficient of determination (R2) values associated with the linear regression model were very low indicating that the model had high residual errors. Also, the p-values associated with the F-test statistic exceeded α = 0.05 significance level for both the M 280 and M 290 machines.

Also, the p-values associated with the F-test statistic exceeded α = 0.05 significance level for both the M 280 and M 290 machines. Therefore, the linear regression model did not provide a statistically significant fit that was better than a constant model and it can be concluded that dimensional accuracy is not linearly related to part location.Table 1: Regression analysis results for dimensional accuracy as a function of part locationMachine M 280 M 290b 0.003 -0.101m 1.38 x 10-4 2.71 x 10-4R2 0.01 0.04F-value 0.45 1.42p-value 0.56 0.24EOS machines require the completion of a “fine-tuning” AM build (see Figure 9) forevery processing parameter set (material type and layer thickness). The fine-tuning build is used for calibration of the beam offset, X scaling factor, and Y scaling factor. These three values are used to account for the melt pool geometry and material shrinkage during laser sintering.

These three values are used to account for the melt pool geometry and material shrinkage during laser sintering. The factors are recognized as an approximation, since the actual melt pool geometry and material shrinkage are dependent upon part size, geometry, and distortion. As a potential solution to the calibration factors’ variation, EOS recommends calibrating the machine using actual parts when possible. In many instances, this is impractical as the AM process lends itself to produce one-off parts and implement frequent part-to-part design changes. Therefore, measurements and calibrations calculated on one build would not necessarily apply to the next build.2037Figure 9: EOS recommended calibration artifact created during "fine-tuning" build.

Measurements collected from the artifact are used to generate beam offset, X scaling, and Y scaling values.Before the start of this research, both machines manufactured a fine-tuning build,measured the resulting geometry, and generated calibration factors for use on all AM builds thereafter. A regression analysis was performed to estimate the correlation between dimensional accuracy and nominal part size, assessing the effectiveness and accuracy of the calculated beam offset, X scaling, and Y scaling factors.The dimensional accuracy data was first plotted as a function of nominal part size (seeFigure 10) with both the machine type, M 280 and M 290, and the measurement direction, X and Y, analyzed separately. Then, both linear and quadratic regression models were fit to the dimensional accuracy data.

In all cases, a quadratic relationship more closely represents the actual dimensional accuracy data. As a result,2039the quadratic fit produces lower residual errors and therefore a larger coefficient of determination. Also, only the linear regression model on the M 280 machine in the Y direction produced a p-value of less than α = 0.05 significance level for the F-test. Therefore, a linear regression model does provide a better fit to the data than a constant model in that case, but it does not provide a better fit in the other three cases. In contrast, the quadratic regression model produces p-values of less than α = 0.05 significance level on all but one case: the M 280 machine in the Y direction. That specific case does not correlate dimensional accuracy to nominal part size at a 95% confidence level.

That specific case does not correlate dimensional accuracy to nominal part size at a 95% confidence level. The other three cases demonstrate statistically significant correlations: the dimensional accuracy data exhibits a quadratic relationship to nominal part size.Table 2: Regression analysis results for dimensional accuracy as a function of nominal part size.

One case, the M 280 machine in the Y direction, did exhibit a linear correlation between dimensional accuracy and nominal part size. Therefore, the linear regression model could be used to calculate new beam offset and Y scaling values to increase accuracy for this particular scenario. However, correcting the scaling factor in only one direction of the build area is impractical, as AM parts require accuracy in both X and Y directions.Ultimately, a more robust technique must be developed and implemented to correct the non-linear errors seen. A better galvanometer calibration technique [8], application of material shrinkage factors on a part-to-part basis, or a pre-processing algorithm to scale parts based on their size could all potentially improve the dimensional accuracy.Conclusions and Future WorkThe requirements of additively manufactured parts are driven by each individual application.

For some applications where dimensional control is not critical or the parts require post-processing such as finish machining, the dimensional accuracy presented in this work is adequate. However, other applications will require tighter dimensional control or the parts will2040p-value 0.49 0.01 0.30 0.30p-value 0.08 0.01 0.02 0.04undergo more cost efficient post-processing such as batch finishing. To meet the design intent in these applications, the existing dimensional accuracy control as presented in this work will require improvements. [7]To assess the root cause of existing dimensional accuracy error in the two EOS machines tested, the build number, machine type, powder state, and manufacturing steps were all analyzed as influencing factors. The statistical results indicated there was no correlation between dimensional accuracy and build number. The machine type however indicated the largest dimensional accuracy variation at 0.088 mm.

The statistical results indicated there was no correlation between dimensional accuracy and build number. The machine type however indicated the largest dimensional accuracy variation at 0.088 mm. Since each machine undergoes its own calibration and material shrinkage factor calculations, there is a lot of variability that could cause this difference. The third factor, powder state, did not exhibit any statistically significant effect on dimensional accuracy. Finally, the manufacturing step showed no statistically significant difference between each subsequent step, but did reveal a statistically significant difference from the very beginning of the manufacturing process (post build) to the end (post part removal). The overall difference was 0.032 mm.

The overall difference was 0.032 mm. Since measurement device’s precision uncertainty was calculated to be 0.010 mm, collecting more data with a more accurate measurement device is desired to confirm the results presents.The effect of part location and nominal part size on dimensional accuracy were alsoanalyzed using a regression analysis. There was no correlation between part location and dimensional accuracy; the data varied randomly across multiple part location distances. The dimensional accuracy data did exhibit a strong quadratic relationship to nominal part size. The non-linear errors resulting from such a relationship make them difficult to correct with the current machine calibration strategy.

The non-linear errors resulting from such a relationship make them difficult to correct with the current machine calibration strategy. The root cause is unknown at this time, however, improved galvanometer calibration, implementation of the material shrinkage factors on a part-to-part basis, or predictive modeling and compensation of the shrinkage prior to build could all correct the error.In continuation of this research, it would be advantageous to independently verify theindividual machines’ accuracy, including the galvanometer calibration. The machine qualification could build upon existing efforts to create a standard test artifact [9] by using the artifact for machine accuracy assessment and calibration. Also, collecting and analyzing additional data from more AM builds will improve the accuracy of the statistical tests and regression models generated thus far.

Also, collecting and analyzing additional data from more AM builds will improve the accuracy of the statistical tests and regression models generated thus far. This may help to develop an understanding of the underlying root cause affecting dimensional accuracy, a necessity to improve the results. The inclusion of more and varied part geometries should also be considered on new AM build layouts. Nominal part sizes should be selected to fill in the existing data’s gaps and also extend range of measurement size past the current maximum. Parts of varying cross sectional area are also important to investigate as they could potentially reveal new trends that would need to be characterized. In parallel to the actual AM builds and build measurements, thermo-mechanical modeling of the AM build layout should be conducted.

In parallel to the actual AM builds and build measurements, thermo-mechanical modeling of the AM build layout should be conducted. The models may be able to predict dimensional accuracy changes and could be used in the future to improve dimensional control before an AM build.2041AcknowledgementsThis work would not have been possible without the contributions of several individuals. We would like to thank Naval Air Systems Command personnel Dr. Venkat Manivannan, Mr. John Schmelzle, and Mr. Eric Kline for their contributions. We would also like to thank Mr. Michael Geyer, Ms. Gabrielle Gunderman, and Mr. Mason Jacoby from ARL Penn State for their efforts designing, measuring, and recording data.This material is based upon work supported by the Naval Air Systems Command (NAVAIR) under Contract No. N00024-12-D-6404, Delivery Order 0321.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDIMENSIONAL ANALYSIS OF METAL POWDER INFUSED FILAMENT - LOW COST METAL 3D PRINTINGShane Terry1 – Ismail Fidan, PhD2 – Khalid Tantawi, PhD31Primary Author, Department of Mechanical Engineering, Tennessee Technological University, Cookeville, TN 38505, SMTerry42@students.tntech.edu2Corresponding Author, Department of Manufacturing and Engineering Technology, Tennessee Technological University, Cookeville, TN 38505, IFidan@tntech.edu3Department of Career Readiness-Mechatronics, Motlow State Community College, Smyrna, TN 37167, KTantawi@mscc.eduABSTRACTThe process of Additive Manufacturing (AM) is the newest form of fabrication with the primary method being layer-by-layer production.

The most common form of this technology is Fused Filament Fabrication (FFF), where material is deposited layer by layer to produce a highly customized part. When compared to subtractive manufacturing the production of waste is greatly reduced. This study presents some innovations on a new metal fabrication technique for FFF printing. By printing a PolyLactic Acid (PLA) compliant metal powder composite filament, a part can be made with approximately 90% metal composition and sintered. The sintering process removes the PLA bonding leaving a 100% metal part fabricated on a low cost FFF printer. Overall, this study reports the initial findings on dimensional changes in low cost metal 3D Printing process.INTRODUCTIONDuring the process of AM, material is deposited additively to produce a part with minimal waste. There exist many methods of AM such as resin-based photopolymer Stereolithography (SLA), Powder Bed Fusion (PBF), and Fused Filament Fabrication (FFF) [1].

There exist many methods of AM such as resin-based photopolymer Stereolithography (SLA), Powder Bed Fusion (PBF), and Fused Filament Fabrication (FFF) [1]. The process of FFF is the safest, least expensive, and most common method of AM. This technique of 3D printing involves the fusion of material in a layer-by-layer process as demonstrated in Figure 1 [2]. This unique form of fabrication has allowed for opportunities in biological [3], concrete[4], and outer space printing [5]533zX• •Figure 1. FFF 3D printing platformThe most common material utilized for this process is polymers such as PolyLactic Acid (PLA). The research presented in this paper is that of the process of Metal FFF (MFFF) printing a field that has been scarcely researched in comparison to other metal 3D printing techniques. By extruding metal polymer composite filament via FFF, parts with approximately 90 vol% metal composition can be fabricated.

MFFF using MPLA material534The sintering procedure is the process in which post-printed parts are heated to near- melting temperatures for the metal powders within the composite material. This results in the fusion of the metal powder, and the removal of the polymer bonding agent. This procedure is common in manufacturing techniques that utilize powdered metallurgy to achieve superior part complexity. Newer metal AM technologies often utilize this method of post-printing processing to yield fully formed metal objects. For the method of MFFF, sintering is utilized to both bond the metal powder and remove the polymer. Specimens at each step in the procedure are further shown in Figure 3.Metal/Pl.A ~rt •fter Sintenf'IIMetaVPl.A part after post p,ocessingFigure 3. MFFF specimens at each stage of fabricationThis method of metal 3D printing provides many benefits when compared to other techniques.

MFFF specimens at each stage of fabricationThis method of metal 3D printing provides many benefits when compared to other techniques. Common methods of metal printing are Selective Laser Sintering (SLS), Selective Laser Melting (SLM), and Direct Metal Deposition (DMD). These fabrication technologies have higher health concerns, are more expensive, and require large areas to operate [6].REVIEWThe overall process of metal 3D printing has been widely explored for methods of fabrication, mechanical properties, microstructural behavior, and useful applications. Metal AM is widely used in fabrication of complex or high part count assemblies most commonly found in aerospace.

Metal AM is widely used in fabrication of complex or high part count assemblies most commonly found in aerospace. The part consolidation and increased part complexity that the metal AM process introduces is irreplaceable by any other method.As discussed in William Frazier’s article on Metal 3D Printing (M3DP) [6], the common M3DP techniques contain processing defects that result in a loss of stress that the fabricated parts can withstand. These processes are also shown to be expensive in both time and money expenditures. The business case assessment will ultimately determine the success of AM as it is currently favored in smaller production lots with the higher cost of raw materials being offset by lower fixed costs related to conventional manufacturing. But there is a large value to be put on AM due to the versatility and adaptability of the process.

But there is a large value to be put on AM due to the versatility and adaptability of the process. Overall, AM is projected to lead to new advancements in manufacturing components and will continue to grow with technology [6].Two of the most common M3DP techniques (SLM and DMD) were examined and compared to other Powder Metallurgy (PM) techniques by a research group out of Nanyang535Technological University [7]. It was found in a comparative study that SLM is a more applicable choice for fabricated parts that require complex geometries such as internal structures and cavitations. The process of DMD is able to produce larger parts at the expense of dimensional accuracy and can serve as an alternative to the fabrication of large-scale components.The current techniques that are able to utilize FFF technology for fabricating metal parts is indirect metal fabrication.

Through the process of casting, patterns and sand-formed parts can be indirectly formed using 3D printed parts that are used to generate a mold [8], [9]. These techniques encompass all of the expenses and safety requirements associated with casting to fabricate parts. A research group from the company Pentair produced a literary review on many forms of M3DP techniques with the benefits and drawbacks of each method being explored [10]. This overview paper gives design and general information for metal fabrication methods. The number of methods that are mentioned and discussed here is very extensive. The group performed a comparison of the various processes in relation to material, labor, and energy costs.

The number of methods that are mentioned and discussed here is very extensive. The group performed a comparison of the various processes in relation to material, labor, and energy costs. Despite being a detail exploration into the various metal print technologies, there were no M3DP techniques that utilized FFF.The research group from Fraunhofer IFAM produced their own metal infused polymer composite material by infusing a Polyamide Nylon (PA) matrix with a metal powder filling [11]. The weight percent of metal polymer composition is varied between 0-65%. The created filament has properties brittle properties and is to remove the bonding agent. The process shown is the only available research that explores the MFFF process.The research group from The University of Texas at El Paso found a method of infusing ABS with a metal powder to form a metal polymer composite material[12]. The combination involves the infusion of ABS with copper and iron powder.

The combination involves the infusion of ABS with copper and iron powder. The metal content is varied from 10- 50% by weight. However, the material is similar despite the research never exploring the post printing sintering process that produces an entirely metal part. The results of this study showed that tensile strength decreased with the introduction of metal powders with an increase in thermal conductivity.PROCEDUREThe MPLA filament that is used for the MFFF process is a metal powder polymer composite material with the highest available metal composition. The printing parameters that are used for the study are based on the combined properties of the composite material. MPLA can be printed with minimal changes to the printing parameters from the base matrix polymer. The settings that need to be modified are 100% infill, decrease in printing speed (around 20 mm/s), and an increase in material flow (around 110%).

The settings that need to be modified are 100% infill, decrease in printing speed (around 20 mm/s), and an increase in material flow (around 110%). Printing parameters can be further tuned with settings such as retraction and fan speed; however, the three aforementioned settings have the most influence on the printability of MPLA.Post-printing, the formed parts must be fired in a furnace/kiln to yield a 100% metal part. This procedure involves the use of a crucible, refractory ballast to support the specimen, and a furnace capable of 1200 C with ramping temperatures. A ceramic crucible is used to suspend the specimen and refractory ballast with minimum contaminants. The two ballasts that are tested in this study are a carbon graphite mixture called Magic Black Powder (MBP) and green sand. The selection of these two materials are based on the preservation of oxidation and dimensional accuracy.

The selection of these two materials are based on the preservation of oxidation and dimensional accuracy. For the temperature and times, the sintering temperature sweeps are shown in Table 1.536Table 1. Description of the time and temperature sweeps used for the sintering processTemperature (°C)Time150Hold for 75 Mins400Ramp over 200 Mins at 1.25 C/Min983Ramp over 180 Mins at 3.24 C/Min983Hold for 240 MinsIn post-sintering, the metal parts are processed to removed oxidation and contaminants. This step is to accurately report the dimensional changes associated with producing a metal polymer composite material, and the yield of the specific metal originally used.

This step is to accurately report the dimensional changes associated with producing a metal polymer composite material, and the yield of the specific metal originally used. For example, this study specifically analyzes the copper polymer composite filament, and therefore the copper oxide remaining in the part post sintering is not included.RESULTSThe process of powder metallurgy allows for highly complex and customized parts however the sintering portion of fabrication involves the removal of the powder bonding agent. The removal of approximately 10 vol% polymer material results in dimensional losses. If the shrinkage associated with the MFFF process is characterized, then during the design stage parts can be sized to compensate.By using cubes that are symmetric across all major cartesian axes and specifically analyzing the copper MPLA, the shrinkage can be better understood.

The two printing parameters that have the greatest influence on precision for this MPLA filament is layer height and print speed. The specimens that are used for analysis are definite with the nomenclature shown in Table 2 in accordance with the changed variables parameters.Table 2. Description of the changed variables for Dimensional AnalysisCU.0225.20.0 MaterialCUCopper FilamentLayer Height03 0225 0150.3 mm 0.225 mm 0.15 mmPrint Speed10 2010 mm/s 20 mm/sSpecimen Size0 10.5 in3 1.0 in3For the sintering process, two methods of firing the specimens is performed. The MBP ballast used is mixed with water in approximately a 1.8:1 ratio mixture; the specimens can be537coated and suspended to reduce oxidation during the sintering procedure. Post-sintering the containments and oxides are removed to obtain a valid value for the dimensional changes. The second sintering criteria utilized the same MBP mixture, time, and temperature sweep.

Post-sintering the containments and oxides are removed to obtain a valid value for the dimensional changes. The second sintering criteria utilized the same MBP mixture, time, and temperature sweep. The changed variable is the ballast in which the specimens are suspended in. For the secondary case, casting green sand is utilized to suspense the specimens. To reduce the oxidation, the parts are coated in the MBP mixture and then pressed into the sand. The working principal behind this is the greater density of the sand will assist in the conservation of dimensional accuracy while the MBP will minimize oxidation.The first sintering criteria resulted in an average dimensional loss due to the MFFF process is approximately 15.8% when averaged between specimens and planar view. The trend of shrinkage shows that the more precise printing parameters result in less dimensional loss.

The trend of shrinkage shows that the more precise printing parameters result in less dimensional loss. The second sintering criteria resulted in an average dimensional loss of approximately 11.1% when averaged between specimens and planar view. Similar to the first process, the trend of shrinkage shows that more precise printing setting result in less dimensional loss.Table 3.

5 0 40 . 5 0 30 . 4 9 80.5 0 6S p e c i m e nA n a l y s i sb yC U . 0 1 5 . 1 0 . 0C U . 0 1 5 . 2 0 . 0C U . 0 2 2 5 . 1 0 . 0C U . 0 2 2 5 . 2 0 . 0C U . 0 3 . 1 0 . 0C U 0 3..2 00D m e n s i o n a liT a b l e 4 .D m e n s i o n a lil o s s e sf r o m u s i n gg r e e n s a n dUsing a vacuum furnace, VF expects an average shrinkage value of 7%. This study is on the process of low-cost metal 3D printing and therefore an open-air furnace is used and the results are within accountable tolerances. The lowest shrinkage was achieved with a green sand ballast and precise printing parameters resulting in a dimensional loss of 5.16%.CONCLUSIONThe process of MFFF is an incredibly new and innovative method of producing metal 3D printed parts. This technology could provide more energy efficient methods of producing metal objects.

This technology could provide more energy efficient methods of producing metal objects. Characterization of this fabrication technique is an important metric for utilization and dimensional characteristic are necessary especially during the design stage. With a minimum dimensional loss of approximately 5%, the printing and sintering procedures could be tuned to acceptable levels for a specific application. These results solidify MFFF as a possible method of M3DP for future industrial and academic applications. Further work will examine the mechanical and microstructural properties to better account for the expected behavior of MFFF fabricated parts.ACKNOWLEDGEMENTSThis study has been made available and funded through NSF Award 1801120, Smart Manufacturing for America's Revolutionizing Technological Transformation, and partially NSF Award 1601587, AM-WATCH: Additive Manufacturing - Workforce Advancement Training Coalition and Hub.

The testing for this research was performed at Tennessee Technological University’s Material Science and Cryogenics Laboratories. The authors of this paper would like to show appreciation to the operators and technicians of these spaces.REFERENCES[1][2]T. Wohlers, Wohlers Report 2019: 3D Printing and Additive Manufacturing State of the Industry. 2019. J. Comb, W. R. Priedeman, and P. W. Turley, “FDM technology process improvements,” Proc. Solid Free. Fabr. Symp., pp. 42–49, 1994.[3] Q.-J. LIANG and X.-D. LI, “Application of FDM Additive Manufacturing Technology in[4][5]Making Medical Surgical Model,” DEStech Trans. Eng. Technol. Res., no. ecame, pp. 133–137, 2018. F. Bos, R. Wolfs, Z. Ahmed, and T. Salet, “Additive manufacturing of concrete in construction: potentials and challenges of 3D concrete printing,” Virtual Phys. Prototyp., vol. 11, no. 3, pp. 209–225, Jul. 2016. T. Prater, M. J. Werkheiser, F. Ledbetter, and K.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDIMENSIONAL COMPARISON OF A COLD SPRAY ADDITIVE MANUFACTURING SIMULATION TOOLR.A. Murray*, G. Foy*, L. Clemon†Advanced Manufacturing Alliance, Charles Darwin University, Darwin, 0909, Australia †School of Mechanical and Mechatronic Engineering, University of Technology Sydney, Ultimo, 2007, AustraliaAbstractHigh-velocity particle spray greatly increases metal additive manufacturing deposition speed over other commercial methods. Accurate prediction and measurement of this process will improve process control. A LightSPEE3D machine fabricated symmetric copper components. On-board software predicts the build geometry (.stl) given the input geometry and the build settings. Assessment of prediction accuracy is needed to enable rapid part design and print setting optimization.

Assessment of prediction accuracy is needed to enable rapid part design and print setting optimization. White-light 3D-scanning and high-fidelity optical microscopy scans are compared to the simulation and intended 20mm cubes using hausdorf distance: 1. Control-repeated scans: 0.38±0.48mm, max:2.25mm 2. Intended-original vs. scans: 1.42±1.58mm, max:6.72mm 3. Software-predicted vs. scans: 0.44±0.66mm, max:3.97mm Discrepancies up to 6.72mm and asymmetric fabrication artifacts were identified. The reduction in the hausdorf distance for simulation vs intended-original, and larger distance of the simulation compared to control, indicate the simulation tool may enable rapid optimization given over/under spray quantification.

Recommendations for reducing asymmetric fabrication artifacts and over/underspray are provided.Keywords: Cold spray, Scanning, Part Inspection, Metal Additive, High-velocity Particle Spray1333IntroductionA key challenge in additive manufacturing (AM) is the speed of material deposition for a given precision. For metal AM, post-processing is often necessary for final parts. High velocity particle spray, sometimes referred to as cold spray [1], embraces post-machining as a necessary step and greatly increases the deposition rate metal AM with this consideration. High-velocity particle cold spray greatly increases metal additive manufacturing deposition speed over other commercial methods. Commercial cold spray machines can achieve deposition rates up to 100g/min [2][Spee3D] and 750g/min [3][Titomic]. To achieve such high deposition rates, precision in each deposited track of material is traded.

To achieve such high deposition rates, precision in each deposited track of material is traded. However, accurate prediction and quantification of this trade-off is not well documented. Post-machining relies on excess built material to achieve desired surface finish and dimensional accuracy for a fielded product. Better quantification and prediction of the overbuilt, or potentially underbuilt, areas is needed. Not all small and medium sized enterprises (SMEs) that could benefit from rapid fabrication offered by this technique maintain a metrology laboratory. Lower cost scanning technologies and open source software may offer SMEs access to fully utilize high-velocity cold spray. Part inspection using 3D scanning has been shown to be speedy, reliable and precise [4]–[7].

Part inspection using 3D scanning has been shown to be speedy, reliable and precise [4]–[7]. This manuscript approaches this challenge by establishing a systematic accounting of errors from solid modelling to fabricated part, comparing lower cost scanning techniques with high-end precision measurements, and reporting the deviations at each step. This quantification of errors and error build-up will help improve prediction and measurement of this process and inform future control algorithms.MethodologyThis research quantifies the error in high velocity particle spray additive manufacturing for a LightSPEE3D machine from initial CAD model to 3D scanned part model. Five stages are identified in the process of creating a component model and producing a scan of the final product. These five stages are identified as:1. Initial solid model (CAD) – the original design of the component prior to adjustment considering machine and process limitations.2.

These five stages are identified as:1. Initial solid model (CAD) – the original design of the component prior to adjustment considering machine and process limitations.2. Print optimised model – the massaged model accounting for known design rules associated with the fabrication machine and process.3. Machine predicted part model – a simulated/predicted outcome of the part generated by a proprietary software native to the machine, Genesys Simulation for this work.4. Fabricated Component – the physical object created using the selected process and machine.5. Scanned part model – a 3D representation of the physical component created using a tactile, optical, or other scanning method.A solid 20 mm cube was designed in CAD, designated ‘initial solid model’. This initial solid model was adjusted for printability on a LightSPEE3D machine by rounding the corners using known fabrication constraints of the machine, designated ‘print optimised’.

This initial solid model was adjusted for printability on a LightSPEE3D machine by rounding the corners using known fabrication constraints of the machine, designated ‘print optimised’. The print optimised model was loaded into the LightSPEE3D machine for fabrication preparation. On-board proprietary software (Genesys Simulation at the time of the experiments, now called TwinSPEE3D) predicts the build geometry (.stl) given the input geometry and the build settings. This software converted the print optimised model into a ‘predicted model’ based on the fabrication settings and output a1334“tool path” for the printer to fabricate the part. Next, the part was fabricated to form a solid copper component affixed to an aluminium plate. A second component was fabricated using the exact same initial model, print optimised model, predicted model, tool path and an identical aluminium build plate and orientation.

A second component was fabricated using the exact same initial model, print optimised model, predicted model, tool path and an identical aluminium build plate and orientation. These components were then scanned using both a Solutionix C500 structured-light 3D-scanner scanner and high-fidelity optical microscopy (Nikon LC15Dx laser scanning coordinate measuring machine). These scanning techniques generate point clouds of data. The point clouds of scans from different orientations are stitched together using software built into each scanning systems. For the Soluntionix, the software is ezScan 2017. For the Nikon, the scanning software is Geomagic. Point stitched and repaired point clouds are then compared in MeshLab [8]. Point clouds are reconstructed using the Poisson surface reconstruction tool [9]. Meshes are aligned using the Align tool, point-based alignment. Then the meshes are frozen in place and Hausdorff distance between them is computed [10].

Meshes are aligned using the Align tool, point-based alignment. Then the meshes are frozen in place and Hausdorff distance between them is computed [10]. The results of the Hausdorff distance computation are reported and used in the error quantification. Colorized images are created by using the Colorize by vertex quality tool in MeshLab. The above mesh alignment and comparison process is used for each scan comparison. The following comparisons are computed:1. Print optimised solid model vs. Scanned Part Model 1 2. Software-predicted vs. Scanned Part Model 1 3. Solutionix C500 scan vs. Nikon LC15Dx scanFor this initial investigation, the simplest model possible has been generated, a unit square. This component allows for initial fundamental examination of errors as they change over the 5 stages of production.

This component allows for initial fundamental examination of errors as they change over the 5 stages of production. The discrepancy between each stage of production is of interest as errors may compound at later stages.Table 1: Stages of fabricationModelFile typesInitial solid model (CAD).stl, .prt, .stpPrint Optimised (CAD) .stl, .prt, .stpSoftware- predicted Model .stl, .otFabricated Partn/aScanned Part Model.stl, .ply, pointsWe measure the difference between electronic representations of the component at each stage. For each comparison we designate a reference and a target model.Discrepancy between the CAD file and an achievable component using this machine is computed using the CAD model as a reference and comparing the predicted model from the native machine tool controller.

Discrepancy between the predicted achievable component from the onboard controller and the actual component is computed.1335ResultsThe original CAD model and software prediction for both components is identical and depicted in Table 1 above. The actual components and scans are slightly different. Visual inspection shows a clear asymmetric artifact in component one on the top surface (Table 2), as well as numerical comparison (Figure 1). This asymmetric fabrication is less obvious in component 2, indicating a statistical variation in fabricated components for the same machine inputs. Two scanning methods are depicted, a high-end high-resolution scan using a Nikon LC15Dx laser scanning coordinate measuring machine. The grey colored sections of the overlayed components indicate areas where the reference scan (Nikon generated) protrudes. The color indicates distance of difference between the two scanning techniques.

The color indicates distance of difference between the two scanning techniques. Cooler colors going toward blue represent less deviation, where warmer colors going toward red indicate a higher deviation.Table 2: Comparison of fabricated components using different scannersPart1Part2PictureHigh-end scanner (Nikon)Low-end scanner (Solutionix C500)1336High-end vs low- end scan comparisonScans of each block were repeated 5 times using the Solutionix C500. Repeated scans show a consistent deviation between the two scanning methods, where the maximum deviation is 0.316 mm. This deviation is significantly lower than the deviation generated by other steps in the fabrication process. The results of comparing the low-end scanner to the reference scans (Nikon) are summarized in Table 3.Table 3: Comparison of low-end scanning with reference scan.

Third largest deviation comes from the repeatability of the machine to make identical rough parts (Figure 1). Deviation between the two tested scanning techniques was the smallest of all deviations.ConclusionsThe results indicate a deviation in scan size using an inexpensive scanning system when compared to a high resolution scanner. This is positive toward enabling SMEs to utilize this process for a subset of work without need for a full metrology laboratory and high-end equipment. The results provide an order of areas that should be addressed to improve prediction and execution of this process. While the highest deviation occurs between the initial solid model and the scanned part model, this is expected as the high-velocity spray process intentionally overbuilds parts so they can be machined later. Effort could focus on reducing this overbuild to the minimum possible, but underspray should be strictly avoided.

Effort could focus on reducing this overbuild to the minimum possible, but underspray should be strictly avoided. The error in the software-predicted model represents an opportunity for improvement. While on average, this deviation is less than 1mm, the maximum deviation is significant at almost 4 mm, which occurs near the base of the components. The base of the components is likely to be consumed in post-machining. This deviation for simple component, such as the 20 mm cube here, a 4 mm deviation is 20% error in one direction. Improving the software prediction for final part geometry is highlighted as a need going forward. More complex components may also change the overspray, or underspray, areas as the spray angle varies.We identified a breakdown of the stages of fabrication and inspection for high velocity particle spray additive manufacturing. The difference in scanning methods for a cost effective scanner and a high end scanner for a simple part were measured.

The difference in scanning methods for a cost effective scanner and a high end scanner for a simple part were measured. The differences indicate potential for adoption of low cost scanning methods for AM parts in this process. The results provide a prioritization of focus areas for future improvement in each of the identified 5 stages, both the software prediction and desired final part could be improved. Discrepancies up to 6.72 mm and asymmetric fabrication artifacts were identified. The reduction in the Hausdorff distance for1338simulation vs initial-original, and larger distance of the simulation compared to control, indicate the simulation tool may enable rapid optimization given over/under spray quantification. Recommendations for reducing asymmetric fabrication artifacts and over/underspray are provided.AcknowledgmentsThe authors would like to acknowledge the technical staff at the University of Technology Sydney for their support.References[1] A. Sova, S.

mensional Issues in StereolithographyDavid L. Winmill, Daniel M. Hoopes and Suresh S. Jayanthi DuPont Somos®, Two Penn's Way, New Castle, DE 19720-2407AbstractNew stereolithography photopolymers have recently been introduced that provide a wider range of functional properties similar to those of high-density polyethylene. One of the important criteria for these materials is the dimensional accuracy and stability in end-use applications as mold masters or the actual functional parts. This work investigates the dimensional stability of one of these new materials with varying amounts of exposure during build. The effect of aging on the part dimensions is reported.

This work investigates the dimensional stability of one of these new materials with varying amounts of exposure during build. The effect of aging on the part dimensions is reported. The result of environmental humidity extremes at ambient temperature on part dimensions is investigated and compared for parts made from two different families of stereolithography resins, namely DuPont Somos® 7100 and Somos® 8100.IntroductionRapid prototyping technologies have gone through several advances in the last few years. Stereolithography in particular has seen some dramatic improvements due to equipment changes and innovations in photopolymers. The primary driving forces for new photopolymers are the quest for improved mechanical properties and dimensional stability.Prototypes made by stereolithography now offer high resistance to harsh environmental conditions.

For example, some of the new materials now provide heat deflection temperatures over 100c C and can also be used in water applications without an appreciable drop in physical to properties. More recent stereolithography materials are providing properties similar commonly used plastics like high-density polyethylene and ABS. All of this has significantly increased the range of applications in which stereolithography can now be used.To achieve a more enduring growth of these applications, a better appreciation of the changes that can occur in the properties of these materials is needed.

The variations in the physical properties with different fabrication strategies, aging under different environmental conditions and endurance testing are becoming more important as these materials are being used to validate not only the design but also the functional intent of the prototypes.This paper focuses primarily on aspects of the dimensional behavior of these new classes of materials and also highlights some of the technological limitations that we have to live with. In trying to build dimensionally accurate and stable prototypes many issues have to be monitored. These include the nature of the geometry and the subsequent slice files as well as the orientation of the part on the build platform.

These include the nature of the geometry and the subsequent slice files as well as the orientation of the part on the build platform. The aspects that are given more attention here are the stereolithography machine calibration, the photopolymer resin being used, the fabrication303style used to build parts, conditions that they experience.their subsequent post-curing and eventually the environmentalThe experimental. studies reported in this paper. are based on two high performance materials for stereolithography. One of these is a high strength, high heat deflection temperature and humidity tolerant material .called Somos® 7110. The other is a high flexibility and high impaqt strength material called Somos® 8100 or 8110 that will be commercialized in the third quarter of this year. This work will first highlight the contribution of machine calibration as a sourqe of error in the production of stereolithography parts.

This work will first highlight the contribution of machine calibration as a sourqe of error in the production of stereolithography parts. Secondly, it characterizes the influence of laser exposure and post-exposure on dimensional stability. Thirdly, it shows the influence of environmental conditions on the dimensional stability of finished parts. Finally, the findings are summarized and observations are made with the view to improving the dimensional behavior.Recent WorkIn 1997, Paul Jacobs (1) first showed that processes such as stereolithography exhibit non-uniform shrinkage behavior and contain a random noise shrinkage component. Following up on Jacobs' work, Tom Mueller (2) expanded the concepts into a more general model that would include the effects of pattern accuracy as well as the shrinkage of injection-molded plastic.

Following up on Jacobs' work, Tom Mueller (2) expanded the concepts into a more general model that would include the effects of pattern accuracy as well as the shrinkage of injection-molded plastic. This would allow tolerances to be predicted for the molded parts rather than the tool.Prior to these two works, significant effort was focused on quantifying the achievable accuracy levels using an industry recognized "User-Part" (3). This particular diagnostic has been It is also used to optimize the compensation factors necessary to improve achievable accuracy. used as a confirmation diagnostic to assess the realistic tolerances achievable in a typical stereolithography machine after all compensation factors have been applied. Figure 1 illustrates the distribution of error across the part for a compensated diagnostic part. As the figure shows there is a high degree (± 6 mil) of random noise.

Figure 1 illustrates the distribution of error across the part for a compensated diagnostic part. As the figure shows there is a high degree (± 6 mil) of random noise. Contributing to the noise are the machine calibration status, process characteristics, possible non-uniform shrinkage behavior of the stereolithography material and the measurement techniques.0.00750.00250.00250.00750.00.40.61.02.12A2J4~4.64.95~6.77.17.485899.195Figure 1: Distribution of Random Noise in a Compensated SL Diagnostic Part (Measured Error on Y axis; CAD Dimensions on X axis)304System Status and. Dimensional AccuracyIn order to investigate the system calibration status and the effect of different exposures on part dimensions, Somos 8100® was used in a 3D System's SLA™ - 500 stereolithography machine. Sixteen square boxes (2 inch by 2 inch) were designed and spaced evenly on the platform in a four by four array.

Sixteen square boxes (2 inch by 2 inch) were designed and spaced evenly on the platform in a four by four array. The CAD for the boxes featured sidewalls as well as an internal circular boss and flat panel. Each of the four columns of boxes received a different hatch over cure of -2, 0, 2, or 4 mils whereas all the other build conditions stayed the same. All of these were quite normal, such as 6 mil layers, ACES build style (4 mil hatch spacing), and a box hatch. The boxes were built without any material shrinkage compensation. A beam width compensation of 6 mils was used.

The boxes were built without any material shrinkage compensation. A beam width compensation of 6 mils was used. Just after the parts were imaged the X and Y dimensions of the boxes were measured and found to be substantially different from the chosen CAD values.Initial Part Size Variation over the SLA - 500 Platform (Part minus CAD Dimension in mils)Row Number 1 2 3 41 7 6 6 7Column Number2 10 9 5 83 5 3 5 44 2 3 3 2Table 1: Dimensional differences for the X dimensions of the two-inch square boxesAs shown above in Table 1, the measured X dimensions were all larger than the desired CAD dimension, and the largest errors tended to occur on the left-hand side of the platform. In the case of the Y dimensions (see Table 2 below), all the dimensions were smaller than intended and the greatest errors occurred on the right hand side of the platform. Clearly then, this machine has a poorly calibrated platform.

Clearly then, this machine has a poorly calibrated platform. Therefore the initial green state dimensions were taken as the starting point from which to measure any changes that might occur due to aging or post-cure procedure.305Initial Part Size Variation over the SLA - 500 Platform (Part minus CAD Dimension in mils)Row Number 1 2 3 41 -12 -12 -12 -11Column Number2 -10 -11 -12 -83 -14 -15 -13 -134 -14 -15 -16 -15Table 2: Dimensional differences for the Y dimensions of the two-inch square boxesIn the future it would seem to be a good idea if the platform calibration was measured with a photo responsive film such as Dylux™ before imaging the parts.

The values obtained from this could then be used as the initial values and would be completely free of any effects (expected to be small directly after a build) of the photopolymer resin.Dimensional Accuracy after ImagingAfter measuring the initial dimensions as described above, two boxes from each hatch over-cure exposure were post-cured with UV light for one hour in a 3D Systems' Post Cure Apparatus. A little later the extents, the circular boss and the panel thickness were measured. The results for the extents (see Figure 2) show that the lowest exposures (-2.0 mils hatch over cure) actually produce an expansion of the parts.4.003.00len«-I 2.00 z- o!!. -w 1.00 ~"wz :E« 0.00 -:r: °0 -1.00.,.,.

-w 1.00 ~"wz :E« 0.00 -:r: °0 -1.00.,.,. .(),,,,X& y- uv I I·.0- • - X & Y - 1 DG I2.004.02.00.02.04.06.0HATCH OVERCURE (MILS)Figure 2: Dimensional Behavior with Different Hatch Over-cures for Somos® 8100The amount of expansion falls with increasing exposure and reaches zero with 4.0 mils hatch over-cure. The boxes from each hatch over-cure that were not post-cured were left for a day at ambient conditions (labeled 1 DO) and again measured. These data are rather scatteredbut show a general trend of the lower exposure to cause shrinkage as the parts age in their green state. The circular boss showed similar effects, but the data was noisier due to the smaller dimension.

The circular boss showed similar effects, but the data was noisier due to the smaller dimension. The panel was too thin (80 mils) to determine any dimensional changes in the Y direction.In order to further investigate the effect of hatch over-cure on dimensional stability in the green state, tensile bars (150 mm long) were measured for four different exposure conditions over a Erolonged period of time after imaging (Figure 3). These parts were made on a DuPont Somos stereolithography machine with 2-mil scan spacing and a raster scan with a spot diameter of about 6 mil. Again the lowest exposures gave the most change, and in the case of parts made with! mil over-cure the dimensional change was + 0.8% after 30 days which equates to 8 mil per inch. However, in the case of higher hatch over-cure exposures of 8 and 9 mil, there was almost no dimensional change with time.0.8c CI: 0_ 0.4 e~ 0-0.2 J:::; c C') .2 i ~ ...J -0.2 ~0.600.4~ .. -.------L-V r? r~·-........__ ..

-.------L-V r? r~·-........__ .. ....-. r~"'" ..... V.... • • .._..IIll. .III. . . . . . . . . . . III..~.-..-----' ~~~~_:~:: I E ----9;11 mil0.60.8o10Time (days)2030 CAD Length =5.9055 in.Figure 3: Dimensional Behavior with Different Hatch Over-cures for Somos® 8100 Green Tensile Bars at Ambient ConditionsEnvironmental Effects and Dimensional StabilityThe effect of the environlnent after the parts were imaged, cleaned .and post-cured was determined on a set of 6 inch tensile bars that had been kept at ambient conditions for about a month. Three sets of six bars made from either Somos® 7110 or Somos® 8110 were sanded lightly, especially on the down-facing surfaces, to remove any remaining supports so that they would be suitable for repetitive weighing. Each set of bars was then left in a large glass jar either at ambient lab temperature (21°C) and humidity, immersed in distilled water,. or at 0 %. relative humidity (RH).

Each set of bars was then left in a large glass jar either at ambient lab temperature (21°C) and humidity, immersed in distilled water,. or at 0 %. relative humidity (RH). At the start and approximately every five days after that each bar was weighed and the length measured.The bars made froll1 Somos® 7110 and held at 0 % RH lost an average of 0.003% weight per day, but their weight was stabilizing. The bars at ambient conditions (Air) actually gained weight slightly presumably due to a slightly increasing humidity in the laboratory with the onset307of summer. The bars in water gained a total of 1.2 % weight over 47 days (an average of 0.03 % Wt per day) and had not yet reached equilibrium (Figure 4).Imaged with 6 mil Layers, 4 mil Hatch Over-cure, UV post-cured~ 1040-:E1.20,2) ~ 1.00 11 0.80 .; ~ 0.60 o E 0.40 ,g 0.20 CD 0) 0.00 Ii tj0.20/' / --- -o./V710~-- ..20~--30--~~~."".

The length of the bars at ambient conditions (Air) varied slightly presumably due to varying humidity in the laboratory. The bars in water gained a total of 0.2 % length over 47 days (an average of 0.004 % per day) but had not yet reached equilibrium.308The bars made from Somos® 8110 and held at 0 % RH lost an average of 0.006 % weight per day, but their weight was stabilizing. The bars at ambient conditions (Air) actually gained weight slightly presumably due to a slightly increasing humidity in the laboratory with the onset of summer. The bars in water gained a total of 2.4 % weight over 48 days (an average of 0.05 Wt% per day) bufhad not yet reached equilibrium (Figure 6).Imaged with 6 mil La~rs, 4. mil Hatch Over-cure, uvpost-cured~ 3.001: 2.50 m ~ 2.00 1i c 1.50 OJ 1:o 1.00Ee 0.50-& 0.00 cot3 -0.50c.7 fA ---.......I-o_ ..-J~./~- - 30- - 30~--------...... - - -.

The bars at 0% >RH lostlengththroughout the study, but then gained some at the end. The length of the bars at ambient conditions (Air) varied slightly presumably due to varying humidity in the laboratory. The bars in water gained a total of 0.6 %length over 48 days (an average of 0.013 Wt % per day) but had not yet reached equilibrium even after that length of time.309ConclusionsTobui1~dill1en~ion~ll¥ia.9CllratePartSb~ steI"eqlithographY'i.machinecalipration isa very important consideration thatshould\ not be overlooked. A way of· measuring the. imaging accuracy before building. parts has been proposed.Ithas been shownthattherandomnoisecomponent of the stereolithography process can a significant contributor •to <the overall part accuracy.

imaging accuracy before building. parts has been proposed.Ithas been shownthattherandomnoisecomponent of the stereolithography process can a significant contributor •to <the overall part accuracy. As more attention is given to the theindividl1~componentsthat<affect dimensional change and therefore random noise, achievableaccuI"acy levels wilLimprove.The amount of exposure used during the part build affects the dimensional stability of the part after building. In general, higher exposure provides more stable parts, whether they remain in the green state or if they are post-cured with UV radiation in the conventional way. This has been shown to be the case with Somos® 8100, one member of a family of stereolithography resins that produce flexible, high impact strength parts.

This has been shown to be the case with Somos® 8100, one member of a family of stereolithography resins that produce flexible, high impact strength parts. The effect of exposure on other stereolithography resins will be the subject of future studies.Afterpartshave.been built and post-cured it has been shown that environmental conditions,in particularhulllidity,. have a pronounced effect on the dimensional stability. This occurred for parts made frQm both.of thestereolithography resins tested, namely Somos® 7110 (a general pUrPose, high strength, ... highheatdefleption temperature resin) ... andSolllos®•.. 8110, although the later resin changed the. most. .For six.,inch tensile bars in some environments, equilibration had notoccurred.over a 47-day period. It was shown that parts left at ambient room conditions.are susceptible to dimensional changes due to humidity changes.

It was shown that parts left at ambient room conditions.are susceptible to dimensional changes due to humidity changes. Further studies· are underway to look into these issues at a greater depth and will be reported in the future.References1.2.3.Paul F. Jacobs, "The Effect of Shrinkage Variations on Rapid Tooling Accuracy" presented at the 4th Annual Eugene C. Gwaltney Manufacturing Symposium, RPMI, Georgia Institute of Technology, Qctober 1-2, J~97. Tom Mueller, "A Model to Predict Tolerances in Parts Molded in Pattern Based Alternative Tooling" presented atthe RapidPrototyping and Manufacturing Conference, May 19-21, 1998. Edward G. Gargiulo and Suresh S. Stereolithography" presented at Prototyping Symposium organized by NIST, October 16-17, 1997."Current State of Accuracy in the NIeasurements and Standards Issues in RapidJayanthi,AcknowledgementsThis work has been made possible by the active participation of some of our valued customers in our technical studies.

We would also like to acknowledge the help we obtained from our colleagues within the Somos® Group and from other divisions within DuPont.310.

mensional Variability AnalysislniPost-Processing Of Rapid ToolingWanlong Wang, James G Conley and Henry W Stoll Department of Mechanical Engineering Northwestern University Evanston, IL 60208AbstractRapid tooling for sand casting can be made by creating pattern with elements made by solid freeform fabrication (SFF) devices. Using this approach, post-processing and hand finishing remain as necessary steps to improve the surface finish quality of the pattern. For rapid·.tooling using laminated object manufacturing (LOM) models, post-processing includes decubing, sanding and sealing followed by integration with a match plate and/or conventional cope & drag pattern elements. Since the critical finishing operations are intensively manual, it is difficult to estimate the dimensional capability of rapid tooling by. LaM process.

Since the critical finishing operations are intensively manual, it is difficult to estimate the dimensional capability of rapid tooling by. LaM process. The Objective of this paper is to use statistics to evaluate dimensional variability associated with postprocessing using the accepted industry best practice.1,,0 INTRODUCTIONSolid freeform fabrication (SFF) refers to the physical modeling of component or tooling geometry using layered manufacturing technologies. These technologies make it possible to quickly generate polymer, wax, or paper--based prototype parts from solid model computer-aided design (CAD) representations. Parts are typically generated by building up one layer at a. time with the thickness of each layer determining the accuracy of the part and the time requiredto the use of SFF build it.

Parts are typically generated by building up one layer at a. time with the thickness of each layer determining the accuracy of the part and the time requiredto the use of SFF build it. technologies logically extends to rapid fabrication of tooling for casting processes such as investment casting, die casting, and sand casting.Initially used for the production of parts for design validation,Over the past decade, rapid tooling (RT) based on SFF models has found. an increased number of industrial applications. For example, the QuickCast system, developed by 3D Systems Inc. [1], is already used in the automobile product development in Ford Motor Company. Gustafson [2] uses laminated object manufacturing (LaM) models as elements of patterns for sand casting.For rapid tooling with LOM, the dimensional capability of the tool making process is a key factor influencing industrial. applications.

applications. While rapid tooling with SFFcanhelp to realize significant time and cost savings, the dimensional capability and stability is not as competitive as that achieved with computer numerical control (CNC) machined pattern production. To understand the dimensional accuracy and consistency of rapid tooling using SFF methods, some research417Wang,. et aI, [31 studied the error sources in the models. Hopkinson, et aI, [4] investigated the thermal..,.nn·.....n1~Asources. We VV. .....v\,JlVJ.J. using theLOM process. The r \ . ."'.nAc''' .......... is then addressed. Since the operations the dimensional loss or increment in the statistics to analyze the dimensional paper. is on rapid tooling using the LaM.... ..,tJ ...·......' ..... V'L"" to otherrapid t9pling processes.Sand casting is an oldianciwicielyused.metalJormingprocess .in which parts are produced by castiJ:l.gJnoltenmetal into sand molds.

..,tJ ...·......' ..... V'L"" to otherrapid t9pling processes.Sand casting is an oldianciwicielyused.metalJormingprocess .in which parts are produced by castiJ:l.gJnoltenmetal into sand molds. It is estimated that more than 90% metal parts are produced by this process/IS].. To. improve the dimensional accuracy, the error sources must be identified. The error sources in the.sandcastinga,reiHustrated in Figure 1 [6].machining. Manual machining is a time-consuming and high cost method and needs experience, knowledge and skills. With this experience-dependent method, it is difficult to guarantee the dimensional accuracy and improve productivity. CNCmachining is a good candidate to improve the efficiency and processing accuracy. <It is. being widely used in today's pattern making industry.

CNCmachining is a good candidate to improve the efficiency and processing accuracy. <It is. being widely used in today's pattern making industry. Figure· 2 gives a flowchart of product development steps for a sand casting process using lartlinatedobject manufacturing (LaM) pattern elements.3.0POST..PROCESSINGERROR SOURCES IN·LOM ··RAPID TOOLINGThe LaM process starts with a 3D computer-aided design (CAD) model of the desired tool or pattern. The model is then tessellated into triangular facets and sliced into layers each having the thickness of a sheet of paper. The LaM model is then constructed by laser cutting and gluiu.g sheets of paper together. Surround material/and material in regions of the part that are hollow mustbe removed in a subsequent "decubing" process. The surface of the resulting LaM model nlust then be sanded to smooth the "stair step" texture created by the layers of paper.

The surface of the resulting LaM model nlust then be sanded to smooth the "stair step" texture created by the layers of paper. Finally, the surface · must be sealed with lacquer or an· equivalent· sealant to keep moisture from being absorbed intdthe model. Each step in this process introduces error as shown in Figure 3.Of the sources identified in ·the.·i figure 1, decubing error and· post-finishing error are those associated with post-processing that can cause delamination or dimensional swelling.

Since this operation is typically conducted manually, it is difficult to make an accurate quantitative analysis for the dimensional change before and after this process, since each would be subjective.418MOLDING METHODSMOLDING EQUIPMENTCore makingPatternsMolding sandOverall casting geometryDIMENSIONAL VARIABILITYMOLDING MATERIALSGEOMETRYFig 1 Fishbone diagram of major factors affecting dimen.sional variability of casting2D Drawing3D CAD solid model& manufacturing ~ (add draft andrequirementshrinkage if needed)Tool design f---+ (Patt.ern design and H core box design)Tessellation(.STL format)Conventional -+ LOM production ~ (Decubing, ~ sand castingPostcprocessingFirst articlesmooth and seal)processPartsFig 2 The sand casting process using rapid t()oling (LOM)Building orientationProcess parametersDistortion in decubingModel Tessellation----;J.T ~~Il....-_ _~_~\__':'>~\l..." Pattern7'productionSanding and sprayShrinkage and expansionFig 3 Fishbone diagram of error sources associated with the LOM process3.1 Decubing Error Residual stress can develop in the LOM build as the result of the temperatllrechange.During the building process, the temperature of the workpiece will rise.

because of heatllsed to iITlprove the adhesive bonding between paper layers. After the building process is finished,. residual stress due419I-C-to volume shrinkage develop within the model as it cools to room temperature. When the model is "decubed", constraints are released and the residual stresses are redistributed. This can result in ~hape distortion of the model and in some cases, delamination of layers. As with other sources of errorin the LOM process, the dimensionalerrortha,tmay occur depends ion the detailed geometry of the part, the part orientation with respe.ct to build direction, paper thickness and other process factors. In general,distortion is greatest in thin wall parts [3].3.2<PO$t-Finishing Error Afterdecuping, the model is usually sanded and sealed. Sanding is necessary to remove "stair step" irregularities and to smooth part features. The sealed lacquer coating seals the surface and strengthe~s the model.

Sanding is necessary to remove "stair step" irregularities and to smooth part features. The sealed lacquer coating seals the surface and strengthe~s the model. The lacquer coating is also sanded to smooth the surface finish of the modeLiThese operations are usually performed. manually by experienced operators. Error introduced during these manual procedures depends on operator skill and specific part geometry. Because of the· number of hard to control factors involved, post finishing error is likely vary significantly with tool geometry.4..0 EX.PERIMENTAL INVESTIGATION4. 1 Testpart An experimental investigation was conducted to statistically analyze the errors discussed above. The test part selected is a thin wall aluminum sand casting used in an aerospace application. A 3D CAD· representation of the part is shown in Figure 4. This part has many typical features including cylindrical surfaces and transitions, planes, and small through holes.

A 3D CAD· representation of the part is shown in Figure 4. This part has many typical features including cylindrical surfaces and transitions, planes, and small through holes. Because of its thin walls, the part is sensitive to decubing and post finishing. Although this is a tooling master, not a casting model, it suits. the purpose of our study. Our purpose is to determine the influence of post-processing on part dimensions. Generally the part is supplemented and extended in some dimensions to add the machining stock).The models was built with the chord height of 0.001 inch. The LOM models were built using a LOM 1015 Plus machine.

The LOM models were built using a LOM 1015 Plus machine. Tessellation and process parameters and resulting build time are given as follows:Laser power: 43 W Laser beam radius: 0.0050 in Plotter speed: 10.0 in/s Heater-slow-speed: 3.00 in/s Heater-fast-speed: 5.00 in/s Platform-slow-speed: 0.50 in/s Paper thickness: 0.0044 in Chord length: 0.001 in Build height: 4.68194 in Total layers: 1089 Elapsed time: 12 hrs 50'4204.2. Dimensional Measurement To assess the effect of tessellation, the dimensions of a variety of features were measured before and after post-finishing using a Brown & Sharpe coordinate-measuring machine (CMM) [7]and compared with the CAD model nominal dimensions. Best metrology practices were used.through out the investigation to ensure that measurement system error did not influence results.

Best metrology practices were used.through out the investigation to ensure that measurement system error did not influence results. Each measurement was repeated five times using the CMMmachine.For each trial, the touch points were randomly chosen and distributed aroundthe surface or plane to insure that aU variation was included. Figure 5 identifies the location of each dimension.Fig 4 The solid model of the test partTo assess the measurement validity and reduce any influence from the measurement error, gage repeatability and reproducibility ( Gage R&R) [6, 8, 9] has been implemented on the CMM machine.4.3 Decubing Error Measurement As discussed above, decubing error may occur due to residual stresses and delamination. To measure the distortion caused by the decubing operation, the concepts of flatness and roundness are introduced. Flatness is the condition of a surface having all. points· in one plane (generally measuring 4 points).

Flatness is the condition of a surface having all. points· in one plane (generally measuring 4 points). Roundness is the condition where all points are in a circle (generally measuring 4 points). It is computed by subtracting the maximum point from the minimum point about the best fit circle [7, 10]. The flatness and roundness· measured here are the flange. plane and the right side of the elbow part (dimensions 13-16, ·see Figure 5). Each measurement has been repeated five times andthe average.value and standard deviation calculated.4.4 Post-Finishing Error Measurement The prototype is sanded and sealed to achieve a final finish according to the conventional post finishing process. This is done by sanding-sealing-sanding-sealing iterations. Sanding andsealing are generally done manually. As such, it is difficult to consistently controLdimensionsofthepart. In this study, the sand paper and lacquer are used to ·sand. and seal.

Sanding andsealing are generally done manually. As such, it is difficult to consistently controLdimensionsofthepart. In this study, the sand paper and lacquer are used to ·sand. and seal. After post-finishing, the prototype can be used as the pattern in sand casting process. Then the •same measurement can be performed on the pattern.421[5]R..22 [8][6][14][15]1][12][16]Figpart dimensions ·l11easured by CMMComparing the items.and values in Table( 1 and 2 and refer to Figure 6,.." 10, we can make the following observations:Generally speaking, the variability of dimensions is the same before post-finishing as it is after post-processing..Of 16 dimensions, there are 6 dimensions greater in standard deviation before than after post-processing, 7 less an<.1.3equal.

As such, there are no obvious trends demonstrating that the standard deviation increased or decreased with geometry.Considering the standard error, there are 4 dimensions greater before than after, 8 less and 4 equal. This suggests that the standard error worsens as a result of post-processing.422For the range, there are 8 dimensions greater before than after, 6 dimensions less and 2 dimensions equal.

For example, for For distance and thickness dimensions, there are 3 dimensions greater before than after, 2 less and 1 equal in standard deviation. For the outer cylindriGal surface dimensions, 1 greater and 3 less. For the inner cylindrical surfa.ce, 1 equal and 1 greater. Fot flatness, 1 less and 1 equal. For roundness, 1 greater and 1 less.4.6 Discussion From the above results, we present the following discussions to further clarify our findings:Post-processing can be useful and flexible to guarantee the dimensional accuracy. fu comparison with CNC finish machining, complement dimensional capability with its sanding-sealing"sanding-sealing process (e.g. overcut area). CNC finish machining isless forgiving if an overcut occurs.

overcut area). CNC finish machining isless forgiving if an overcut occurs. While comparing with manual making, the overall dimensional accuracy can be achieved and the productivity is relatively better.post-processingcanPost-processing can repair or reduce some systematic error sources inherited from previous process stages, i.e. paper thickness, tessellation,. build direction, etc. As indicated by Wang, et al[3],systematic errors are not avoidable because ofthe nature of layered manufacturing. But post-processing can help to repair or reduce these errors by its sanding and sealing iterations, if the operator has a better understanding of the process and is experienced with mold making.Different rapid tooling processes may be more suitable to different geometry. According to our experience, the LOM process is good for block, thick wall and very complicated surfaces geometry, but not for small holes or bars, thin walls,<shallow surfaces and difficult undercuts.

Dimension 8 (diameter is 0.22 inch) has verified this. The error is evident.even though the chord height is already then1inimum. Other ••SFF processes (e.g. SLA,FDM)may be better suited to handle the geometry feat\1res that challenge LOM.4255.0 CONCLUSIONFrom the above analyses, we may findthatther~ are no obvious relations of dimensions existingconsistency or the accllrate relationship. Butif we•. can. understand this and make full use of the advantag.es.of post-processing, the. systematic. errors· maybe repaired or reduced and thus the dimensionalcapabilityluaybe improved. GeQmetrydrivers for LOMprocess are also discussed. The <flexibility inpost--processingcan be helpful for rapid.tooling in processing accuracy and thus strengthenthe pattern matches competitive position in the sand cast tool making market.Acknowledgmentl'hison-going research is supported bythe American Metalcasting Consortium/Defense Logistic Agency (AMC/DLA).Lead Time and Cost Reduction Program.

We are indebted to Milos Coric who made and decubed the LOMmodels, and performed the post-finishing.References1. K.· Denton, P. Jacobs, "Quickcast<& Rapid Prototyping: a case history in Ford Motor Company", The 5th International Conference on Rapid Prototyping, Dayton, Ohio, June12-15, 1994, p301-320. R.Gustafson, Guinn, ' - ' ........, ••.o.J.J.I"".. v 85 n1995. p 48~50.Wang,ConleyandH. W. Stoll, "Rapid Tooling Error Analysis for Sand Casting",Transactions of theAFS 102ndCastingCongress, Atlanta, GA, May 10-13, 1998.4. N.Hopkinson andP.Dickens, "Thermal effects on accuracy in the 3D Keltool process", Proceedings of Solid Freeform Symposium, The University of Texas at Austin, Texas, August.11-13, 1997. p 267-274.5. S.S.>Pak, D.A. Klosterman, et al,"Prototypetooling andlow volume manufacturing through International Conference on Rapid laminated object·· manufacturing (LOM), The.· 7th Prototyping, San Francisco, CA, March 31-April4, 1997, p325-3316.E.Peters, R.C.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDirect Additive Subtractive Hybrid Manufacturing (DASH) – An Out of Envelope MethodMatthew C. Frank1, Ola Harrysson2, Richard A. Wysk2, Niechen Chen1, Harshad Srinivasan2, Guangyu Hou1, Carter Keough21 Industrial and Manufacturing Systems Engineering, Iowa State University of Science and Technology Ames, Iowa, USA 2Edward P. Fitts Department of Industrial and Systems Engineering, North Carolina State University Raleigh, North Carolina, USAAbstract:This work addresses a critical issue affecting the adoption of metal additive manufacturing (AM) methodologies; creating a system that can produce a mechanical product to final geometric specification. To date, almost all functional metal AM parts have required secondary processing that in many cases can more than double the cost of the final part.

To date, almost all functional metal AM parts have required secondary processing that in many cases can more than double the cost of the final part. A Direct Additive and Subtractive Hybrid (DASH) manufacturing system using both additive and then subtractive processing has been developed so that mechanical parts can be “digitally manufactured” to meet the final required geometric accuracy. The approach includes the development of a software system to link additive and subtractive manufacturing, using extensions to the AMF format, to maintain product design features along with their tolerance specifications. It also introduces the idea of sacrificial fixtures that are automatically designed into the parts to allow subsequent fixturing in the CNC mill. Once in the milling machine, a part localization system identifies the part and its location, along with capturing the geometry of any remaining AM support material left on the part.

Once in the milling machine, a part localization system identifies the part and its location, along with capturing the geometry of any remaining AM support material left on the part. Finally, all CNC code is automatically generated and the finishing process can be executed in a lights-out operation. This work provides a drastic reduction in post processing time and cost. It further enables expansion of metal AM and uniquely addresses the challenge of out- of-envelope hybrid manufactured parts.Introduction:For many years, the emergence of additive manufacturing (AM) technologies was viewed as the basis of the elusive push-button manufacturing process, and there was a renewed optimism with the arrival of functionalized metal processes such as Electron Beam Melting (EBM) and Selective Laser Melting (SLM). Unfortunately, the high precision specifications of today’s engineered components have not yet been achieved with metal additive processes.

Unfortunately, the high precision specifications of today’s engineered components have not yet been achieved with metal additive processes. In most cases for metal parts, removal processes like machining (milling, grinding and polishing techniques) are inevitable to attain functional part accuracy and surface finish. In fact, most users of metal AM will regularly deploy some form of post process machining for at least a few critical features if not entire sections of critical exterior geometry. This is analogous to the day-to-day operations of the metal casting industry, and therefore, one could argue that it has been received with some undue acceptance. However, the problem with using conventional machining methods for finishing AM parts is that the economies of scale available in the metal casting industry can be used to justify the investment; whereas in AM, we need the capability to effectively produce just one or a few parts as desired.

Unfortunately, there has been no plug-in method of integrating a part from an additive process into a subtractive process directly without significant human intervention (fixturing and orientation set- up). More importantly, process can require several days and in most cases, several weeks1853particularly in the case of sophisticated contours and high precision components (which are typical of commercial aerospace and biomedical parts). For small batch production an efficient means of meeting the geometric and surface finish requirements would decrease overall AM part costs and increase competitiveness of AM processes.Overview of Method:This work presents a method of integrating AM and subtractive operations into a sequential hybrid manufacturing process called Direct Additive and Subtractive Hybrid (DASH).

The DASH system for use in the manufacturing of metal products can combine any direct metal additive manufacturing process with a 4-axis CNC milling system to produce high quality and high precision metal parts with limited human intervention. The system is highly software driven, using the AMF file format to represent and optimize the design for finishing by adding fixture elements and machining allowance prior to AM fabrication.

The system is highly software driven, using the AMF file format to represent and optimize the design for finishing by adding fixture elements and machining allowance prior to AM fabrication. This work includes development of a new software system, integration and enhancements of multiple existing technologies, optimization of a sacrificial fixturing system, and assembly into a unified operational system (Figure1).(a)(b)(c)(d)(e)(i)(h)(g)(f)Figure1: DASH Process Flow; (a) Customer model input, (b) AMF model creation with critical features identified, (c) Machining setup orientation planning, (d), Addition of machining allowance and sacrificial fixture supports, (e) Metal AM processing, (f) clamping in CNC mill, (g) scanning and part localization, (h) automated toolpath generation, (i) finished part after support removalFigure 1 shows a specific example part that underwent these steps and demonstrates the current capabilities of the DASH process.

To start, the customer model (Figure1a) is used to generate an AMF file which is marked up digitally to identify critical part surfaces (Figure1b) in a custom program called AMFCreator (Srinivasan et al., 2015). This step is where the user is choosing surfaces that will require post processing, and providing data about the surface type and tolerance. Next, we take a forward-looking step and determine how the part will need to be machined, using that information to drive a CNC setup orientation (Figure 1c). The AMFCreator software then adds machining allowance to critical surfaces, along with sacrificial fixture elements that will1854secure the part during machining (Figure 1d). Finally, we reach the typical starting point in metal AM, where we have an STL file ready to process and print (Figure 1e). Upon removal from the build plate and AM support removal, the part is clamped in a CNC machine using the sacrificial fixture supports (Figure 1f) (Boonsuk and Frank, 2009).

Upon removal from the build plate and AM support removal, the part is clamped in a CNC machine using the sacrificial fixture supports (Figure 1f) (Boonsuk and Frank, 2009). Next, the part is scanned (using laser or other) and the part orientation and stock conditions are captured (Figure 1g). In CAM, using CNC- RP (Frank et al., 2004), the part has automated toolpaths generated and executed (Figure 1h), and the machined metal AM part is delivered (Figure 1i). The following sections provide a brief review of each major step in the DASH process.Critical feature identification:As a part of the DASH process, one must first identify the critical features that are slated for post process machining. As such, a software package called AMFCreator’ was developed to allow an operator to create and manipulate AMF Files that designate as-printed versus as-machined features.

As such, a software package called AMFCreator’ was developed to allow an operator to create and manipulate AMF Files that designate as-printed versus as-machined features. The software is used across numerous steps such as; 1) importing and exporting to and from STL and PLY files, 2) Selecting surfaces and associating them with features, 3) adding tolerance information to features, 4) Adding machining allowances, 4) computing feature parameters and finally 6) creating sacrificial support geometry for the final machining process.AMFCreator was written in C++ with the QT (“Qt - Home,” n.d.) widget toolkit and the VTK (“VTK - The Visualization Toolkit,” n.d.) visualization system. XML I/O was performed with the TinyXML2 library (“TinyXML-2,” n.d.). The UI consists of two sections, as seen in Figure 2. To the left is a window in which the AMF model is displayed.

XML I/O was performed with the TinyXML2 library (“TinyXML-2,” n.d.). The UI consists of two sections, as seen in Figure 2. To the left is a window in which the AMF model is displayed. Models displayed in this window canFigure 2: AMFCreator software interface1855be scrolled and panned similar to any CAD package. The window is also used to demarcate surfaces for association with a feature, and as a means selecting for volumes and surfaces. To the display window is a list that displays either the set of features or the set of volumes in the AMF file. In this step of the process, the user navigates across the surface of the part selecting and painting features, along with tolerance and other feature parameters.

In this step of the process, the user navigates across the surface of the part selecting and painting features, along with tolerance and other feature parameters. The output of this is an AMF model that is interpreted by the machining process planning software, CNC-RP, which is described next.theright ofFigure 3: CNC-RP Process, (a) Rotary fixture setup; (b) Process sequence of steps (1- 6) to produce a steel componentMachining Approach:An existing enabling technology for rapid machining, CNC-RP, was originally conceived at Penn State and further developed at Iowa State University (Frank et al, 2004, 2006; Petrzelka and Frank 2010). It is a fully automated Subtractive Rapid Prototyping process that uses a 3-axis vertical milling machine with a 4th axis indexer for multiple setup orientations.

It is a fully automated Subtractive Rapid Prototyping process that uses a 3-axis vertical milling machine with a 4th axis indexer for multiple setup orientations. In this system, round stock material is fixed between two opposing chucks and rotated between operations using the indexer, and visibility analysis of cross sectional slice data provides a basis for automated setup planning about a single axis. This implementation uses a modified Greedy set cover algorithm to determine orientations. For each orientation, all visible surfaces are machined using simple toolpath planning while leaving a structure of “sacrificial supports” that are used to fixture the part (keeping the part attached to the remainder of the stock). The number of rotations required to machine a model is dependent upon its geometric complexity. Once all of the operations are complete, the supports are severed in a final series of operations, and the part is removed.

Once all of the operations are complete, the supports are severed in a final series of operations, and the part is removed. Figure 3 illustrates the process steps for creating a typical complex part using this method. The following sections will overview a few critical steps within CNC-RP process planning, for orientation, setup, and sacrificial fixture planning. As shown, CNC-RP is a plug-in for MasterCAM software, using tool library and toolpath generation methods from the commercial software, while automating decisions and planning normally done by the NC code programmer.Orientation and Setup Planning:The “Axis Analysis” interface in CNC-RP is made up of nine images as shown in Figure 4, where each image denotes an orientation and the scores for this orientation given.

As shown, images # 1, 2, 3 always present the three principle orientations, while images # 4, 5, 6 pick the best option according to a sole criterion, namely, the best orientations for visibility, stock diameter and support rigidity. Image #7 presents the software-suggested orientation based on a multi-criteria search,1856while image #8 presents the orientation that the user chose in the search space. Lastly, image #9 presents a rotatable orientation that the user can The colored customize. progress bar the the color visibility, where the indicates the type of surfaces in question (flats, holes, or freeform).indicatesOnce an orientation for the rotary axis the in creating software aids sacrificial supports that keep the part fixed to parent round stock. The user selects 1) a type, and 2) a material maximum allowable deflection (related to overall part tolerances).

The user selects 1) a type, and 2) a material maximum allowable deflection (related to overall part tolerances). CNC-RP then automatically identifies the diameter, number and location of supports based on beam an optimization routine. The interface shown in Figure 5 allows the user to examine sacrificial support placement based on the orientation to modify the positions and sizes of the supports. is an approximately 2X factor of safety, allowing the user to as-much-as delete one of the supports entirely.is chosen,theoryandthey chose andThereFigure 4: Axis analysis interfaceAllowances and Fixturing:Up to this point along the work flow, part’s machining orientation has been chosen and support geometry is identified. Any changes to part orientation is captured in .XML data that is passed back to AMFCreator.

Any changes to part orientation is captured in .XML data that is passed back to AMFCreator. Next, AMFCreator must add additional fixture clamping elements,theFigure 5: Sacrificial support placement interface1857along with machining allowance (padding material) on the feature that will be surfaces machined. In addition, sacrificial supports defined in CNC-RP as cylinders are converted to prismatic (square section) and oriented to minimize AM supports (based on build orientation). Finally, the part’s STL model is prepared for metal printing. As shown in Figure 6, the customer’s initial model in arbitrary orientation is now appended with fixture and support elements that are ready for printing. AMFCreator will eventually contain a library of interfaces fixture depending on CNC machine setup; in this case, a simple 3 jaw chuck clamp-able set of triangles.

AMFCreator will eventually contain a library of interfaces fixture depending on CNC machine setup; in this case, a simple 3 jaw chuck clamp-able set of triangles. An example part build from an EBM is shown in Figure7.Figure 6: Part model in AMFCreator, with sacrificial supports, fixture clamping elements, and machining allowance on critical surfacesclampingFigure 7: Part, AM supports, and Sacrificial supports to fixture clamping elements after powder bed metal printingScanning Localization:andlike Each part, those shown in Figure 7, will be clamped between opposing jaw chucks of the milling machine. Of course, a combination of metal printing variability, warping, handling, placement and clamping errors have placed the part in an orientation that does not precisely coincideFigure 8: Laser scanning interface for DASH1858with the part model in the CAM software.

We now desire the answer to two questions; 1) where is the part geometry located? and 2) how much material represents the “stock surfaces” after printing, now for machining? With respect to the second question, after printing we might anticipate some AM supports have been removal, but not all; therefore scanning yields the as- printed stock conditions. One would actually anticipate that a DASH-processed part would not technically require any AM support removal per se; but it would reduce machining time if AM supports can nominally be broken away before machining. The method chosen to find the part location and model its stock condition is laser scanning. Figure 8 presents the scanning software interface for DASH. In this process, the point cloud data from scanning is used to do two things.

Figure 8 presents the scanning software interface for DASH. In this process, the point cloud data from scanning is used to do two things. First, we identify a set of critical surfaces on the part, and calculate the transformation matrix that would align this as- fixtured part orientation to the CAD model’s location in the CAM (MasterCAM) environment. As such, the part is translated and rotated in CAM, using the CNC-RP software plug-in. In addition, a slightly oversized STL mesh of the entire part, support, etc. geometry is passed to CNC- RP, where MasterCAM can utilize it as a stock model. In the following step, CNC-RP will have all information necessary to generate NC code for critical surfaces.Figure 9: CNC-RP toolpath planning interfaceToolpath Generation:The details of all surface-based automated toolpath planning is neglected here for lack of space.

In brief, each critical part surface previously identified in the AMF model is used to create a set of feature-specific toolpaths. As shown in Figure 9, a flat surface (red) on a part has been identified and machining angles are chosen automatically (Frank et al., 2006). The subsequent steps choose tools, machining strategy, cut parameters, and automatically define containment boundaries.Figure 10: Roughing toolpath calculation1859Once all critical surfaces identified, CNC-RP are calculates additionally rouging for toolpaths operations that remove all remaining AM supports first (Figure 10), before feature- into launching specific toolpaths. Total process planning time for is NC code generation measured in minutes.Implementation:The DASH process has been executed on sample linkage components using both steel and titanium alloys. The titanium alloy was printed using an Arcam EBM while the steel parts were made on an EOS DMLS system.

The titanium alloy was printed using an Arcam EBM while the steel parts were made on an EOS DMLS system. Critical features were originally overgrown 2mm for the first builds in steel and titanium, and then a second set of trials in titanium used a smaller allowance of 1.25mm with no negative effects. Figure 11 the milling shows machine while Figure 12a shows the progression of the part, from as printed in the EBM, to as-fixtured and machined in a HAAS VF2ss, to final part upon sacrificial support In addition, CMM removal. measurement data (Figure 12b) shows dimensional accuracy well within 0.1mm (0.002”).

measurement data (Figure 12b) shows dimensional accuracy well within 0.1mm (0.002”). These and numerous other trials have shown the effectiveness of the DASH process at integrating Additive and and Subtractive Manufacturing providing high-quality finished metal components.the partinFigure 11: DASH processed part immediately after machining(a)(b)Figure 12: Titanium part sample; (a) progression across the process, and (b) CMM analysis data on nominal dimensions1860Conclusions:The proposed DASH process is focused on integrating additive manufacturing (AM) with a subtractive manufacturing (SM) process for the production of precision metal components. The concept for finishing metal additive manufacturing (AM) parts using a traditional subtractive manufacturing process appears to work well for test components to date and is providing high quality parts at remarkable levels of automated process planning.

However, the process is not a simple “turnkey system” and requires human intervention across the stages. It does not, however, require days or weeks from a machine shop nor does it take the level of skill of a typical CNC programmer. In closure, DASH shows promise for significant reductions in both cycle time and skill required to get a functional, post-processed part from metal additive manufacturing.Reference:Boonsuk, W. and Frank, M.C., “Automated fixturing for a rapid machining system”, Rapid Prototyping Journal Vol. 15, No. 2, pp. 111-125, 2009Frank, M.C., Wysk, R.A., and Joshi, S.B., “Rapid Planning for CNC Machining – A New Approach to Rapid Prototyping”, Journal of Manufacturing Systems, SME, Volume 23, No. 3, pp. 242-255, 2004Frank, M.C., Wysk, R.A., and Joshi, S.B., “Determining Setup Orientations from the Visibility of Slice Geometry for Rapid CNC Machining”, Journal of Manufacturing Science and Engineering, Transactions of the ASME, Vol. 128, No. 1, pp.

The problem has been that conventional engineering solutions used to address repair, durability, and weight, typically result in increased acquisition costs and, consequently, are untenable to the program managers. Further, it has become abundantly clear that the cost, weight, and durability of naval air weapons systems must be improved. For example, $3B per year is spent on corrosion-related maintenance of Navy aircraft, curtailing our ability to acquire new weapons systems and decreasing the availability of operational assets [2].Direct digital manufacturing (DDM) is an innovative part fabrication and repair technology that represents a game changing advance in the way airframes are designed, built, and maintained. DDM technology was identified in a prior Navy workshop as a means of effecting a 30% cost reduction and a 30% increase in through put [3].High value, difficult to machine, and process alloy components materials favor the use of DDM.

These types of alloys are typically used in demanding, fatigue-critical applications. Consequently, producing DDM parts with fatigue properties comparable to wrought products isan important objective. In situ and post fabrication techniques were identified as a means of enhancing fatigue performance. Control of surface roughness was also considered important.Executive Guidance Direct Digital Manufacturing (DDM) aligns well with and supports executive level guidance. The CNO [4] and the Quadrennial Defense Review Report [5] have emphasized reducing the Total Ownership Cost (TOC) of our weapon systems. The NAE S&T Strategic Plan [6] and NAE S&T Objective document [7] include incorporating affordability into platform design and construction; and responsive and visible logistics to enable distributed forces.

The President has provided guidance to pursue transformational solutions and support visionary thinkers proposing high-risk, high-payoff research [8].Vision State The vision for DDM is illustrated in Figure 1 [9]. Its implementation would enhance operational readiness, reduce energy consumption, and reduce total ownership cost by exploiting advanced metallic direct digital manufacturing technologies. Metallic parts would be fabricated on- demand at a location in proximity to their end use.Figure 1DDM Vision State: Parts-on-Demand – “Ship Electrons not Parts”Conceptually, a broken or worn part is identified by a maintainer. The maintainer uses a computer terminal to access a parts database. The part’s DDM build package is sent to a fabrication site (e.g., the Navy’s Fleet Readiness Centers) near where it is needed.

The maintainer uses a computer terminal to access a parts database. The part’s DDM build package is sent to a fabrication site (e.g., the Navy’s Fleet Readiness Centers) near where it is needed. The metallic part is fabricated to net shape using direct digital manufacturing equipment, finished, machined and assembled.In the event a build package is not available, the part may be reverse engineered. The needed part could be laser scanned and imported into a digital cloud map. This information could then be transferred to any number of engineering design CAD/CAM programs, e.g., Pro-E. Once the design configuration is established, a DDM build package would be developed and archived for future use.Workshop Purpose The workshop was held in order to assist ONR and NAVAIR develop a robust research and development program in the area of DDM, e.g., an ONR Future Navy Capabilities Program.

The workshop was structured to help identify research opportunities and to address the technical challenges associated with using DDM of metallic components. The technical obstacles to DDM implementation were examined and approaches to overcoming the identified barriers formulated. The workshop explored (i) innovative design concepts which reduce weight, (ii) maintenance and repair concepts, (iii) life-cycle-costs reductions, and (iv) qualification and certification methodology. Further, the overarching goal is to enhance operational readiness, reduce energy consumption, and enable parts-on-demand manufacturingWorkshop Concept of OperationThe workshop provided an opportunity for seventy-two invited experts from Navy, DoD, industry, academia, and leading research institutes to share their views and make strategic recommendations.

Figure 2 provides a list of represented organizations at the workshop.Figure 2 Organizations Participating in the May 2010 Navy DDM Workshop.Plenary A plenary session was held on the morning of the first day of the workshop. Table 1 lists the plenary speakers representing executives and technical leaders from the government, industry, and academia. This was followed by parallel working group breakout sessions in four topic areas: DDM Science & Technology, Qualification & Certification Methodology, Innovative Structural Design, and Maintenance & Repair. Workshop participants were asked to validate the goals and objectives of the working group and to identify technical challenges and approaches for achieving the goals.Table 1Plenary Speakers and their AffiliationDeputy Commander Fleet Readiness Centers, NAVAIRMr. Garry Newton Mr. Richard Gilpin Mr. Greg KilchensteinMr. Mike DeitchmanMs.

Garry Newton Mr. Richard Gilpin Mr. Greg KilchensteinMr. Mike DeitchmanMs. Karen TamingerDirector, Air Vehicle Engineering Department, NAVAIR Senior Sustainment Technology Policy Analyst, OSD (AT&L) delivered by Constance Philips, National Center for Manufacturing Sciences Deputy Chief of Naval Research, Naval Air Warfare and Weapons Science and Technology Department ONR Senior Materials Research Engineer, PI for Materials and Structures for the Subsonic Fixed Wing Aircraft, NASA Langley Research CenterDr. Thomas Donnellan Associate Director for Materials and Manufacturing, ARL PennMr. Blake Slaughter Prof. Dave BourellState. Metallic Processing Group, Boeing Research and Technology. University of Texas Austin delivered in the workshop brief-out session.GOTChA Process The GOTChA (Goals, Objectives, Technical Challenges, and Approaches) approach was used as a tool to develop the products of the workshop. Figure 3 graphically illustrates the GOTChA process.

Figure 3 graphically illustrates the GOTChA process. The Navy defined the goals and objectives of the workshop. The overarching goal of the workshop was to enhance operational readiness, reduce total ownership cost, and reduce energy consumption by exploiting advanced metallic direct digital manufacturing technologies.Figure 3 Methodology: Goals, Objective, Technical Challenges & Approaches (GOTChA)The working groups met for approximately 8 hours. Their first task was to validate and, if necessary, amend the three working group objectives. The working group then identified and prioritized the technical challenges associated with achieving the objectives. The balance of their time was used to develop viable approaches to solve the challenges identified. These approaches were broadly grouped into three time frames: near (< 5yrs), mid (5-10yrs), and far (>10yrs). The workgroup results were briefed-out in a plenary forum prior to concluding the workshop.

The workgroup results were briefed-out in a plenary forum prior to concluding the workshop. A post workshop analysis was performed in which the salient information was summarized and packaged in a format suitable for dissemination.Workshop ResultsThe results of the workshop are divided into three parts: (i) Plenary, (ii) Working Groups, and (iii) Summary.Plenary Summary The plenary session provided a forum for leaders from the Government, Academia, and Industry to discuss DDM in terms of warfighter needs. There was overwhelming agreement that DDM is an agile and viable source of manufacturing and repair and its exploitation would help address pressing naval aviation needs.DoD-Navy’s Environment: The Navy’s current and, likely, future environment was discussed. The Country is at war and naval aviation must respond quickly and effectively to warfighter needs. There are several key factors which will drive naval aviation to use direct digital manufacturing.

The Country is at war and naval aviation must respond quickly and effectively to warfighter needs. There are several key factors which will drive naval aviation to use direct digital manufacturing. There is an increased emphasis on reducing the cost of Defense Department’s Operation: acquisition and sustainment. The average age of our Navy’s aircraft is 19.18 years . This is putting an increased strain on the Navy’s supply chain. As aircraft age, parts that were never expected to break or fail do [10]. Supply chain does not have the ability to repair or produce new parts. Aircraft are grounded while we spend precious time researching vendors who can repair or replace the parts.Naval Aviation Needs: In order to respond to the Country’s wartime footing and aging fleet of aircraft, there has been an increased demand for one-off parts, crash damage repair, and rapid solutions to Red Stripes [10]. An Agile and Viable Source of Manufacturing and Repair (e.g., DDM) is required.

An Agile and Viable Source of Manufacturing and Repair (e.g., DDM) is required. DDM may be especially useful addressing the issues which arise in the Sun Down/Disposal phase of a weapon systems life cycle.The technical barriers associated with inserting DDM were discussed. Some of the more salient research and development needs are listed in Table 2. In order to achieve the vision state of parts-on-demand, the need to accelerate part qualification and certification tops the list. The value of having the capability of producing an aircraft parts in a matter of hours is of little value if it takes weeks or months of testing and evaluation to certify it for use.

The value of having the capability of producing an aircraft parts in a matter of hours is of little value if it takes weeks or months of testing and evaluation to certify it for use. An accelerated qualification process is closely linked to a number of factors including (i) an understanding of machine-to-machine variability and repeatability, (ii) accurate, predictive process models for microstructure and properties, and (iii) computationally guided processes and closed loop process controls.Table 2 Research and Development Needs: Plenary SessionAccelerated qualification and certification methods • Accurate, predictive process models for microstructure and properties • Fatigue properties comparable to wrought materials, Post fabrication processes to enhance fatigue properties. Methods to reduce surface roughness of partsPart-to-part and machine-to-machine variability and repeatability • Computationally guided processes and closed loop control.

Methods to reduce surface roughness of partsPart-to-part and machine-to-machine variability and repeatability • Computationally guided processes and closed loop control. Integration of sensors into process control systems to enable real-time NDE during processingTechnology fusion, i.e., laser scanning, database, design tools, and database • New structural design & analysis tools, e.g., stiffeners that follow load paths • Alloys designed for DDM fabricationWorking Group Products DDM Science and Technology: The Navy’s S&T Objectives for DDM are listed in Table 3. Associated with each objective, the Working Group identified three technical challenges which must be addressed.

Associated with each objective, the Working Group identified three technical challenges which must be addressed. For example, to obtain fatigue properties equivalent to wrought alloys, the Working Group felt that the top two technical challenges were the control of microstructure and the elimination of defects.Table 3 DDM S&T Objectives and Technical ChallengesGoal: Enhance operational readiness, reduce energy consumption, and reduce total ownership cost by exploiting advanced metallic direct digital manufacturing technologies.Objectives and ChallengesObjective 1: Static and fatigue performance equivalent to wrought1. Controlling Microstructure 2. Elimination of Defects 3. Material ChallengesObjective 2: Achieve Statistically Repeatable and Predictable Processes1. Monitoring and Controlling the Process 2. Need to model the process 3. Need NDE for DDM materialsObjective 3: Surface Finish / Minimize Assembly and Post Deposition Processing1.

Monitoring and Controlling the Process 2. Need to model the process 3. Need NDE for DDM materialsObjective 3: Surface Finish / Minimize Assembly and Post Deposition Processing1. Maintain acceptable deposition rates while achieving good surface finish 2. Need controlled position of DDM machine 3. Lack of scientific/technical info on post treatments of DDM surfacesMaturation of technologies at the TRL 3-4 to TRL6 and above are of significant interest. Analysis of the DDM S&T Roadmap, Figure 4, suggests that the following research areas should be explored. In order to enhance the fatigue performance of DDM parts, develop in situ DDM processes to improve part surface finish and to mechanically work the DDM part during deposition, e.g., by using laser shock peening.

The development of physics based models for structure-property-processing are also essential in order to be able to consistently and accurately produce DDM parts of diverse configurations.Figure 4DDM S&T RoadmapQualification and Certification: The three Navy objectives associated with the qualification and certification of DDM metallic components are presented in Table 4, along with the list of technical challenges developed by the Working Group. For example, the Working Group felt that a DDM database of material properties was necessary. Importantly, the database must represent the pedigree of the data and come from a stabilized DDM process.Table 4 Qualification and Certification Objectives and Technical ChallengesGoal: Enhance operational readiness, reduce energy consumption, and reduce total ownership cost by exploiting advanced metallic direct digital manufacturing technologies.Objectives and ChallengesObjective 1: Qualification of DDM fabrication processes1.

Definition of methods required for verifying control of the key process variables 2. Definition of qualification requirements for each of the three phases of product life cycle - new design/prototype, repair/replacement, and production part3. Demonstrate repeatabilityObjective 2: Eliminate the need to qualify each part individually1. Need materials properties database developed under a “stable” process 2. Determine necessary steps to avoid repeating generation of materials properties database for different DDM technologiesObjective 3: Reduce the time & cost of qualification by 90%1. Definition of acceptance requirements for repair/replacement of fielded parts and prototype2. Prioritization of action items associated with qualification processAnalysis of the Qualification and Certification Roadmap, Figure 5, suggests that the following research areas should be explored. Maturation of technologies at the TRL 3-4 to TRL6 and above are of significant interest.

Maturation of technologies at the TRL 3-4 to TRL6 and above are of significant interest. In order to secure the vision of parts-on-demand, accelerated part qualification methods need to be developed. Research should be focused on the development of heuristic, and probabilistic, qualification methods. The development of industry specifications and standards are necessary as is a robust materials database. However, it is unlikely that S&T funding can be used to directly support their development.Figure 5 Qualification and Certification RoadmapInnovative Structural Design: The objectives and the technical challenges associated with Innovative Structural Design are presented in Table 5. In order to reduce structural weight, better, integrated structural & DDM fabrication design tools are needed.

In order to reduce structural weight, better, integrated structural & DDM fabrication design tools are needed. Similarly, to reduce cost, true net-shape fabrication methods are needed which eliminate the need for secondary (post fabrication) processing.Table 5 Innovative Structural Design Objectives and Technical Challenges Goal : Enhance operational readiness, reduce energy consumption, and reduce total ownership cost by exploiting advanced metallic direct digital manufacturing technologies.Objectives and Challenges Objective 1: Reduce structural weight by 25% with no increase in acquisition cost. (freeform technologies, Ti-6-4, apply to entire air vehicle vs. piece part) 1. Better structural optimization including multifunctional design 2. Integrated design while allowing for maintainability 3. Introduce/integrate new alloy composition “sweet spots” for the design 4.

Better structural optimization including multifunctional design 2. Integrated design while allowing for maintainability 3. Introduce/integrate new alloy composition “sweet spots” for the design 4. Design community to accept paradigm shift to adopt new design rulesObjectives 2: Enable complex part fabrication with a 50% reduction in cost.1. Near-net shape manufacturing (reduce buy:fly ratio, reduce labor cost, reduce # of secondary operations)2. Use of lower cost starting materials 3. Develop more innovative techniques to develop/finish near-net shape parts with the goal using the part as-builtObjectives 3: Reduce the design, engineering, build, test & qualification time cycle by 60%.1. No recognized standards exist 2. Streamline design process 3. DDM design with minimal tooling & facilities 4. Risk tolerance by approving authorities & designersThe examination of the Innovative Structural Design Roadmap, Figure 6, reveals two significant features. 1.

Risk tolerance by approving authorities & designersThe examination of the Innovative Structural Design Roadmap, Figure 6, reveals two significant features. 1. There is a need for integrated structural and materials design optimization tools to support the design and certification of non-traditional design concepts. 2. In line with what the Qualification and Certification working group reported, there is a need for a shared materials property database; and industry specification and standards.Figure 6 Innovative Structural Design RoadmapMaintenance and Repair: The objectives and the technical challenges associated with Maintenance and Repair are presented in Table 6. In order to reduce the time required to acquire out-of-production parts by 90%, the need for an improved and accelerated means of qualifying and certifying DDM parts is needed. The objective of reducing part energy content was associated with the need to reduce or eliminate post fabrication processing.

The objective of reducing part energy content was associated with the need to reduce or eliminate post fabrication processing. Further to effect a reduction in logistics foot print, DDM equipment versatility is needed.Table 6 Maintenance and Repair Objectives and Technical ChallengesGoal: Enhance operational readiness, reduce energy consumption, and reduce total ownership cost by exploiting advanced metallic direct digital manufacturing technologies.Objectives and ChallengesObjective 1: Reduce time to acquireout-of-production parts by 90%1. Certification and approval 2. Merging existing technology & material capabilities with actual parts that are needed.3. Quality control – parts feedstock and processObjective 2: Reduce total energy content by 60%1. Post processing: Surface finish requirements & dimensions. Residual stress/heat treatment2. Equipment efficiency increases 3. Raw material cost/deposition efficiency.Objective 3: Reduce logistic footprint by 20%1.

Residual stress/heat treatment2. Equipment efficiency increases 3. Raw material cost/deposition efficiency.Objective 3: Reduce logistic footprint by 20%1. Equipment versatility 2. Additional post processing/NDI/HIP 3. Raw materials storageThe Innovative Maintenance and Repair Roadmap is presented in, Figure 7. The Working Group identified qualification-by-similarity as a needed research area as well as improved surface finish and dimensional accuracy. These two areas were also identified by other working groups as important areas to pursue. The development of alternatives to hot isostatic processing (HIP) of DDM parts to achieve wrought fatigue properties was identified as a critical need.

The development of alternatives to hot isostatic processing (HIP) of DDM parts to achieve wrought fatigue properties was identified as a critical need. The use of HIP is capital intensive and would increase part cost and fabrication time.Figure 7 Maintenance and Repair RoadmapCritical Research Areas SummaryThe critical areas of needed research may be divided as follows: (i) Underlying DDM Science, (ii) Process Control, (iii) Qualification, and (iv) Innovative Structural Design.Underlying DDM Science A robust understanding of the underlying science of DDM is essential to the widespread implementation of DDM. New alloys must be developed that lend themselves to DDM. A priority is the development of physics based models relating microstructure, properties, and process to performance. This represents the foundation upon which process control, part qualification, and innovative designs concepts can be built.

This represents the foundation upon which process control, part qualification, and innovative designs concepts can be built. Innovative in situ processes (e.g., hybrid laser and electron beam systems) and an improved understanding of structure-property- processing relationships are required in order to enhance fatigue properties. This includes an understanding of how to reduce surface roughness. Lastly, technology integration and fusion is required to achieve the desired “Vision State.” The various component technologies associated with DDM fabrication, reverse engineering, qualification, and design must be made to work together seamlessly.Process Control The ability to achieve the vision state of parts-on-demand requires accurate and predictable control of the DDM fabrication process. Machine-to-machine variability must be understood and controlled.

Machine-to-machine variability must be understood and controlled. Industry specifications and standards for the processing of aerospace alloy components must be developed and inculcated throughout the industry. The highest priority should be given to developing integrated in-process, sensing, monitoring, and control technologies.Qualification The ability to produce a wide variety of parts-on-demand is essential to harnessing the full potential of DDM. Part-by-part certification is costly and time consuming and is antithetical to achieving the Navy’s vision. Therefore, alternatives to conventional qualification methods must be found; these are likely based upon validated models, probabilistic methods, and part similarities. Likewise, industry specifications and standards for DDM and DDM processed aerospace alloys are needed. These standards and specifications form the foundation for rapid part certification.

Likewise, industry specifications and standards for DDM and DDM processed aerospace alloys are needed. These standards and specifications form the foundation for rapid part certification. Similarly, advanced NDE techniques capable of detecting critical flaws and defects with a high degree of certainty are needed.Innovative Structural Design The introduction of new and innovative structural designs requires that the design community be knowledgeable of DDM. A priority should be given to the development of integrated structural and materials design tools. A robust DDM database accessible to these communities must be developed in order to provide pedigree data which integrate structural and materials.AcknowledgementsI’d like to thank all those who contributed to the success of the DDM Workshop including the organizers and the participants. The workshop organizers included Chairpersons Tom Mellies, Robert Kestler, Madan Kittur, Brian Boyette, and, Jeffrey Waldman.

The workshop organizers included Chairpersons Tom Mellies, Robert Kestler, Madan Kittur, Brian Boyette, and, Jeffrey Waldman. The workshop was sponsored by ONR and the Navy Metal Working. Special thanks goes to Malinda Pagett, ONR Code 35, my co-organizer for her hard work and leadership.References1. William E. Frazier, “Navy Workshop Aims to Cut Cost”, ADVANCED MATERIALS &PROCESSES. AUGUST 20082. David A. Forman, Eric F. Herzberg, James C. Tran, Amelia R. Kelly, Paul N. Chang,Norman T. O’Meara, Ph.D., “The Annual Cost of Corrosion for Navy and Marine Corps Aviation Equipment” ( LMI Government Consulting, Report MEC70T3, May 2008). 3. William E. Frazier, “Naval Aviation Materials and Manufacturing Workshop: Accelerating Technology Insertion” (Presented at Joint Technology Exchange Group: “Roadmap to Successful Sustainment Technology Insertion” August 26-28, 2008)4. Admiral G.

Admiral G. Roughead, “CNO’s Guidance for 2010, Executing the Maritime Strategy”September 2009 (http://www.cffc.navy.mil/cnog_2010.pdf)5. Quadrennial Defense Review Report. DoD, United States) February 2010( http://www.dtic.mil/futurejointwarfare/strategic/qdr2010.pdf) (6. Naval Aviation Enterprise Science and Technology Strategic Plan. (NAE: CommanderNaval Air Forces, Commander Naval Air Systems, and Director Air Warfare Division) July 1, 2006.7. Naval Aviation Enterprise Science and Technology Objectives, (NAE: Commander NavalAir Forces, Commander Naval Air Systems, and Director Air Warfare , Deputy Commandant for Aviation) April 2010.8. Peter R. Orszag and John P. Holdren, S&T Priorities for the FY 2011 Budget (White HouseMemorandum, M-09-27, Aug 4, 2009)9. DDM ONR FNC POM-12 Proposal, September 2009. 10. William E.

RECT EXTRUSION FREEFORMING OF CERAMIC PASTESXiaopeng Chi, Hongyi Yang, Shoufeng Yang and Julian R.G. Evans Department of Materials, Queen Mary, University of London, Mile End Road, London, E1 4NS, UK Reviewed, accepted September 14, 2006AbstractMicroextrusion freeforming of ceramic lattices from high solids ceramic pastes provides multi-scale hierarchical void structures with the advantages of low shrinkage stress and high sintered density. Alumina lattices were directly fabricated using 80-500 (cid:80)m diameter filaments. We report here on the implementation of design and fabrication of these scaffolds for band gap materials and micro fluidic devices.Keywords: scaffolds, extrusion freeforming, meta materials.IntroductionSolid freeform fabrication (SFF) is a class of advanced manufacturing technologies which generate accurate geometrical objects directly from a three-dimensional computer image without part-specific tooling or human intervention1.

Many SFF technologies have been developed to date and they vary in terms of building speed, surface finish, dimensional accuracy and material properties2-5. A sub-class of methods is based programmable microextrusion of filaments onto a building platform and includes fused deposition modelling FDM6, fused deposition of ceramics FDC7, multiphase jet solidification8-10 and robocasting5. Another direct write assembly11 method uses pH-controlled gelling of ceramic colloids or polymers to build positive or negative scaffolds12.In these mouldless ceramic fabrication methods, the main difference is the method of state change. Robocasting uses temperature control on a copper substrate to transform each layer from pseudoplastic (61 vol. % powder) to dilatant (64 vol. %)5, 13, 14 for which further deformation is prevented.

Robocasting uses temperature control on a copper substrate to transform each layer from pseudoplastic (61 vol. % powder) to dilatant (64 vol. %)5, 13, 14 for which further deformation is prevented. To avoid this complex control of drying kinetics, colloidal inks have been developed for robocasting and are assembled under non-wetting oil to prevent drying during assembly15, 16.The method of state change used here is decrease of solvent content. The paste, composed of ceramic, polymer and solvent has high yield strength but is able to weld to the previous layer before solidifying by solvent evaporation. Sufficient polymer content ((cid:124)32vol.% based on wet paste) eliminates extrusion defects resulting from dilatancy often found in water-based systems. High ceramic content (minimum based on wet paste(cid:124)48vol.%) results in low sintering shrinkage. The different volume percent based on wet paste and on dry paste is shown in table 1.

High ceramic content (minimum based on wet paste(cid:124)48vol.%) results in low sintering shrinkage. The different volume percent based on wet paste and on dry paste is shown in table 1. Low solvent content results in low drying shrinkage. The high yield strength of the paste minimises deformation during fabrication giving higher manufacturing precision.304Table 1. Different volume percent based on wet paste and on dry pasteVolume percent Based onVolume percent Based onwet paste/vol.%dry paste/vol.%polymer3240Ceramic powder4860solvent200This method knits any conceivable two-dimensional extruded filament pattern layer by layer. The non-volatiles (ceramic + polymer) constitute typically 70-80 vol.% mixture of ceramic powder and polymer and 20-30 vol.% volatile solvent (usually propan-2-ol).

The non-volatiles (ceramic + polymer) constitute typically 70-80 vol.% mixture of ceramic powder and polymer and 20-30 vol.% volatile solvent (usually propan-2-ol). The drying-induced shrinkage after assembly was minimized to reduce distortion of the lattice.Experimental DetailsEquipment A press head driven by micro-stepper motors (50,000 steps/rev, supplied by ACP&D Ltd., Ashton-under-Lyne, UK) with a 64-1 reduction box driving 1mm pitch ball screws (Automation Ltd., Oldham, UK) was used to extrude the paste. Hypodermic syringes (HGB81320 1ml, Hamilton GB, Ltd., Carnforth, U.K.) were used as extrusion plunger, piston and nozzle.

Hypodermic syringes (HGB81320 1ml, Hamilton GB, Ltd., Carnforth, U.K.) were used as extrusion plunger, piston and nozzle. Although shortened hypodermic needles performed well in early experiments on filament diameters greater than 200 (cid:541)m, smaller diameters blocked due to the long land length and orthogonal lead-in.Superior extrusion dies were assembled from water-cut jet nozzles made from thin sapphire plates with smooth orifices, some models having a 45 degree lead-in cone. These were inserted into hypodermic needles. Very fine filaments (50 (cid:541)m) can be extruded from these without blocking. A sensor (Flintec Ltd., Redditch, UK) monitored extrusion load and provided overpressure alarm.The three axis table (Parker Hannifin Automation, Poole, Dorset, UK) comprises linear motor driven XY axes and stepper driven Z axis (Figure 1). Three axis motion, extruder drive, pressure and data recording were integrated by Labview software.

Three axis motion, extruder drive, pressure and data recording were integrated by Labview software. An earlier version with a lead screw table locked at sharp reverse turns and was abandoned.305acbdFigure 1. Extrusion freeforming apparatus, (a) twin extrusion presses on the Z axis, (b) computer control system, (c) load cell displays and alarms (d) linear motor XY table.Paste preparation and extrusion Alumina powder (HPA-0.5 Morgan Matroc Ltd, UK) was used for extrusion. Poly(vinyl butyral) (PVB), grade BN18 (Whacker Chemicals, UK) was used as the binder with additions of a grade of poly(ethylene glycol) (PEG) that is liquid at ambient temperature, (MWt = 600, VWR, UK). The PVB and the PEG were fully dissolved in propan-2-ol (GPR, VWR, UK) in proportion 75 wt.% PVB – 25 wt.% PEG. A powder blend with HA 75 wt. %-TCP 25 wt.

The PVB and the PEG were fully dissolved in propan-2-ol (GPR, VWR, UK) in proportion 75 wt.% PVB – 25 wt.% PEG. A powder blend with HA 75 wt. %-TCP 25 wt. %, 17, was added to provide a ceramic/polymer mixture with 60 vol.% of ceramic based on the dry mass and dispersed by ultrasonic probe (IKA U200S, IKA Labortechnik Staufen, Germany) for 15 minutes. The suspension was agitated on a roller table with zirconia media for 12 hours and dried to a paste for extrusion.Results and DiscussionThe final solids content determines the success of lattice fabrication. Each of five different pastes (30, 36, 39, 45 and 50 vol. % ceramic powder based on wet paste) were extruded to give lattices of different macro-porosity (50, 60, 70 and 80 %). In the fabrication process, the ‘sinking’ effect, whereby filaments sag between base-layer supports, results from high solvent contents and is shown in Figure 2.

In the fabrication process, the ‘sinking’ effect, whereby filaments sag between base-layer supports, results from high solvent contents and is shown in Figure 2. The ‘curling’ effect, in which the upper layer peels away from the base layer due to drying-induced stresses is shown in Figure 3a. The schematic diagram of the ‘curling’ effect is shown in Figure 3b. The angle of deformation, (cid:537), can be obtained by the formula below:306tan (cid:32)(cid:84)H L(1)where, H is the drying deformation height; L is the drying deformation length along thefilaments.Figure 2. Filament ‘sinking’ effect resulting from high solvent content in an unsintered alumina lattice with 39 vol.% ceramic powder in the wet paste. The distance between filaments is 630(cid:541)m.baFigure 3. Drying defect in a 14 layer alumina lattice fabricated by the 39 vol.% ceramic powder in the wet alumina paste.

The distance between filaments is 630(cid:541)m.baFigure 3. Drying defect in a 14 layer alumina lattice fabricated by the 39 vol.% ceramic powder in the wet alumina paste. The distance between filaments is 630(cid:541)m.Two factors influence the ‘sinking’ effect: the yield stress o f the paste which is mainly nced by solids fraction and the distance between filaments (span). Using the same paste, influe the span was systematically decreased which in turn decreases the macroporosity of the lattice18,30719. Figure 4 shows that the largest detectable sinking distance decreases as the span decreases, i.e. as the lattice becomes less porous. Using a lattice with fixed span and hence fixed lattice porosity and systematically varying the solvent content with fixed ceramic-polymer ratio, the sinking distance decreases as the non-volatile fraction increases.Figure 4.

Sinking deformation as a function of the non-volatile fraction of the paste for different lattice spans; the largest observed inter-filament deformation was recorded. Note that a high lattice macroporosity (as given in the key) corresponds to a high span.The curling effect is caused by the drying shrinkage stresses duri ng fabrication. If the paste es a fully dense binary of ceramic and polymer after drying, the volume shrinkage is equal becom to the volume fraction of solvent. The linear shrinkage is 10% for 27 vol.% solvent.

If the paste es a fully dense binary of ceramic and polymer after drying, the volume shrinkage is equal becom to the volume fraction of solvent. The linear shrinkage is 10% for 27 vol.% solvent. Four factors influence the extent of curling and delamination: (i) the powder fraction in the dried ink controls the elastic modulus of the 0-3 composite such that modulus typically increases steeply as the powder fraction exceeds 50 vol.%, (ii) the rate of volatilisation of solvent influences the differential shrinkage between layers; in the worst case scenario the base layer has fully dried so that its shrinkage is complete when the next layer begins to dry, (iii) the span between filaments determines the number of welds per unit area and hence overall laminate strength, (iv) as the span decreases and the lattice becomes more densely packed, the vapour transport within the structure decreases and drying rate slows down, in turn reducing drying deformation.308Figure 5.

Deformation angle as a function of non-volatile fraction of the paste for different lattice spans.From Figure 5, when the non-volatile content was less than 49 vol.%, the rendered layers remain wet while new layers are deposited. The increased weld strength between filaments counteracts delamination. Since the first layer is strongly adherent to the substrate, deformation did not occur whatever the lattice span. When the non-volatile content were between 49 and 65 vol.%, the weld strength decreased as the solvent content increased and could not counteract the shrinkage of filaments. Although the overall shrinkage decreased, the drying time was reduced. In this regime, drying deformation increased with increasing non-volatile content. When the non-volatile content was between 65 and 74 vol.%, the overall shrinkage of filaments decreased and drying deformation decreased with increasing non-volatile content. When the non-volatile content was more than 74 vol.

When the non-volatile content was more than 74 vol. %, the solvent was correspondingly low and so the linear shrinkage was low. The effect of lattice macroporosity is very pronounced. At high porosity, the number of welds per unit area is low and delamination is facilitated. Furthermore, the more open structure allows faster drying of rendered layers.When the non-volatile content was greater than 74 vol. %, all these defect phenomena wereabsent as evidenced by the regular structures shown in Figures 6 and 8.Theoretically, the span of the lattice can be made in any size by increasing the non-volatile volume fraction. In practice, the spans between filaments cannot be increased without limit16, 18. If the solvent content was less that 16 vol.%, the paste was too dry to weld to the base layer. The nozzle was easily blocked and the layers did not join. An ideal lattice was attained when the solvent content was approximately 26 vol. %.309Figure 6.

The nozzle was easily blocked and the layers did not join. An ideal lattice was attained when the solvent content was approximately 26 vol. %.309Figure 6. Unsintered alumina lattice with 26 vol. % solvent and inter-filament distance 630 (cid:541)m showing no sagging or drying deformation.The extrusion distance is the gap between the die exit and the previous layer. If the gap is too large, the filaments are difficult to control and may drag across the substrate. If the distance is set too small, the nozzle smears the paste and damages the filaments. The appropriate distance depends on the nozzle diameter and the viscosity of the paste. For the 150(cid:541)m diameter nozzle, the best distance was 0.2 mm.

The appropriate distance depends on the nozzle diameter and the viscosity of the paste. For the 150(cid:541)m diameter nozzle, the best distance was 0.2 mm. The nozzle is raised on the Z axis before a new layer is started.Theoretically, the filament extrusion velocityrVand the velocity of the XY tableXYVarecompliant with:Vxy(cid:32)(cid:167) (cid:168) (cid:169)R r(cid:183) (cid:184) (cid:185)2VR(2)where R and rare the diameter of the plunger and the nozzle respectively.IfrV<XYV, the filament was stretched and ifrV>XYVit formed curves. In practise, theviscous nature of the paste also delayed the response of the stress to the strain19 so that filamentextrusion speedrVdid not respond quickly to changes in the speed of the XY tableXYV. Somany data around the ratio(cid:167) (cid:168) (cid:169)R r(cid:183) (cid:184) (cid:185)2were tested to find out the optimalrVandXYV.

Somany data around the ratio(cid:167) (cid:168) (cid:169)R r(cid:183) (cid:184) (cid:185)2were tested to find out the optimalrVandXYV. To increasefabrication speed, rapid XY table movementXYVand extrusion velocityrVwere needed, butthis places high loads on the microextrusion press. So in this experiment, the speed of the XY310tableXYVwas 6.735 mm/s and the speed of the filament extrudedrVwas 0.00825 mm/s.The alumina lattices were sintered at 1540(cid:113)C for 1 hr after slow heating to 400(cid:113)C at a rate of 2(cid:113)C/min with 1 hour dwell so that the polymer in the samples could be removed slowly and evenly. The microstructure, observed by SEM and shown in Figure 8 is of near full density, with isolated residual porosity and average grain size in the region 3-4 (cid:541)m.Figure 7.

The microstructure, observed by SEM and shown in Figure 8 is of near full density, with isolated residual porosity and average grain size in the region 3-4 (cid:541)m.Figure 7. Microstructure of alumina filament sintered at 1540(cid:113)C at 5(cid:113)C/min with 1 hour dwell after heating to 400(cid:113)C at 2(cid:113)C/min.Lattice design and fabrication Itthat periodic dielectric structures can block propagation of electromagnetic waves. One such structure is the ‘woodpile’ 20-22. In this approach, woodpiles with different distances between filaments can be made by alumina filaments. Such ‘woodpiles’ with 630 (cid:541)m between filaments are shown in Figure 8 and their dimensions in Table 2.is well knownTable2.

Such ‘woodpiles’ with 630 (cid:541)m between filaments are shown in Figure 8 and their dimensions in Table 2.is well knownTable2. The dimensions of 20 layer alumina lattices with 74 vol.% non-volatile content.Length /mmWidth /mmHeight /mmFilament Diameter /mmGap /mmBeforesintering19.6819.682.900.150.63Aftersintering15.7015.702.300.120.50311Shrinkage /%20.2Figure 8. Unsintered alumina lattices comprising 20 layers prepared from a paste with 26 vol.% solvent, having an inter-filament distance of 630 (cid:541)m.When defects are made periodically in an electromagnetic band gap (EBG) structure, defect mode resonance can be created23, 24. So a lattice with square lacuna was designed and fabricated. The amplified square area of the lattice was drawn by Solidwork software as shown in Figure 9a. It was obvious that the width of the square was limited by the distance between filaments.

The amplified square area of the lattice was drawn by Solidwork software as shown in Figure 9a. It was obvious that the width of the square was limited by the distance between filaments. However the distance between filaments cannot be made too big or filaments will sink. The sample with the square lacuna periodic structure was fabricated by extrusion freeforming and is shown in Figure 9b. The size of the samples and the shrinkage rate is given in Table 3.abFigure 9. (a) Magnified square lacuna area of the lattice; (b) Micro-extruded alumina lattice with square lacuna.Table3.

The size of the samples and the shrinkage rate is given in Table 3.abFigure 9. (a) Magnified square lacuna area of the lattice; (b) Micro-extruded alumina lattice with square lacuna.Table3. Dimensions of lattice with square lacunaLength /mmWidth/ mmHeight /mmFilament Diameter/mmGap/ mmLacuna size/mmShrinkage /%Before sintered After sintered30.6424.6230.6424.622.902.300.150.120.480.381.59x1.591.26x1.2620.7%312As applications for microfluidic reactors broaden, materials with higher temperature capability and chemical resistance are sought to fabricate the microfluidic reactor25. Ceramic microchannel devices have been fabricated by a prototyping process chain combining stereolithography and low-pressure injection molding26, and by lamination27. An alumina microfluidic reactor was fabricated by extrusion freeforming in this work. The reactor was made up of five parts (shown in Figure 11a).

An alumina microfluidic reactor was fabricated by extrusion freeforming in this work. The reactor was made up of five parts (shown in Figure 11a). The upper and lower covers were separately made up each of two layers assembled from alumina filaments attached each other tightly. The flow channels were made up of four layers assembled by alumina filaments with two large channels (2.55mm) to facilitate gas flow and 20 smaller channels (0.45mm) for the gas to disperse well. The mixing chamber was made up of four layers assembled by alumina filaments with 20 small channels (0.45mm) for gas mixing. The outflow was made up of four layers assembled by alumina filaments with one large channel (0.9mm) for the gas to exit. A schematic diagram of the two gas flows is shown in Figure 11b. The overall size of the reactor was 15.48 x 10.68 x 1.92mm after sintering and the small channel for the gas mixing was 0.36mm after sintering.

The overall size of the reactor was 15.48 x 10.68 x 1.92mm after sintering and the small channel for the gas mixing was 0.36mm after sintering. An alumina microfluidic reactor without the lower cover is shown in Figure 12.abFigure 11 a) cutaway view of the alumina microfluidic reactor; b) schematic diagram of the two types of gas through the reactor.Figure 12. The alumina microfluidic reactor without the lower cover313ConclusionsIt has been demonstrated that regular ceramic lattices can be fabricated by extrusion freeforming by adjustment of paste composition. Conditions for avoiding the characteristic defects of the process such as filament sagging between previous layers and drying deformation were obtained by controlling the solid loading of the paste, the extrusion rate, the X, Y table velocity, the filament diameter. Upon sintering, the alumina lattices have sintered densities greater than 95% and sintering shrinkage 20%.

RECT FABRICATION OF METAL ORTHOPEDIC IMPLANTS USING ELECTRON BEAM MELTING TECHNOLOGYO. L. A. Harrysson1, D. R. Cormier1, D. J. Marcellin-Little2, K. R. Jajal11 Department Of Industrial Engineering, North Carolina State University, Raleigh NC 27695,2 Department of Clinical Sciences, North Carolina State University, Raleigh NC 27606[Reviewed, accepted August 13, 2003]AbstractMetal orthopedic implants have been used for many decades with great success. Replacement joints and plates for bone fractures are usually made from titanium, cobalt- chromium or stainless steel. Through recent advancements in biomodeling, custom orthopedic implants can be designed. However, fabrication of these custom implants can be prohibitively expensive with traditional processes. With the introduction of Electron Beam Melting (EBM), direct fabrication of fully dense metal components is possible.

With the introduction of Electron Beam Melting (EBM), direct fabrication of fully dense metal components is possible. In this paper, the development of titanium for the EBM-process will be discussed, and direct fabrication of custom designed orthopedic implants made out of steel and titanium will be demonstrated.IntroductionOver 500,000 Americans undergo joint replacement surgery each year, with hip and knee implants being the most common. The number is predicted to exponentially increase over the next couple of decades1. Many different types of bone plates are used to fixate bones after a severe fracture or after an osteotomy. Joint implants and bone plates are usually made out of stainless steel, titanium or cobalt-chromium and come in standard sizes and generic shapes. Even though total joint replacements have been successfully used for several decades, the longevity of the implant components is not satisfactory in many cases, especially in younger patients [1, 2].

Even though total joint replacements have been successfully used for several decades, the longevity of the implant components is not satisfactory in many cases, especially in younger patients [1, 2]. One of the most common causes for implant revision is loosening of the component, which is mostly caused by bone remodeling [3]. Bone is a living tissue that is constantly changing due to external forces to optimize its structure and minimize its weight [4]. Bone will increase in density when experiencing a dynamic load, and it will decrease in density when experiencing a static load or no load at all. This is a common problem for astronauts when spending an extended amount of time in microgravity. Older people experience the same problem when they become less active, and their condition is referred to as osteoporosis.When implanting a joint replacement, the affected portion of the joint must be removed to provide healthy bone to interact with the implant [5].

Modern implants are coated with a porous titanium coating on the bone-implant interface surface to enhance and promote bone ingrowth. These implants do not use any type of adhesive to secure the implant component to the bone and solely depend on the bone ingrowth to provide a secure and stable attachment. This is the preferred approach to be used on all healthy patients with sufficient bone growth. Due to the current manual surgical tools available to the surgeon, the bone-implant interface on a standard knee implant is designed with straight, flat surfaces (see figure 1.) [6].1 Data provided by American Academy of Orthopaedic Surgeons, Dept. of Research and Scientific Affairs439Figure 1. Computer models of a human distal femur prepared for implantation and a generic femoral implant component.The geometry limitations cause areas with increased pressure and areas with decreased pressure, which causes bone remodeling and finally loosening of the implant components.

To further increase the problems, all humans are different, and standard implants only come in 5-7 sizes with generic shapes. The longevity of the implant is highly dependent on the initial fit and how well it resembles the natural shape of the joint [7]. Hip and knee implants have had the same basic design for several decades, which have been limited by the hand tools available to the surgeons. The manual saws and drills can only produce straight cuts and holes, and the precision is highly dependent on the surgeon’s skill. Recently, two robotic surgical systems have been commercialized that can perform the cutting and drilling operations with far better precision using end mills as actuators [8, 9, 10]. A robot also has the ability to easily machine freeform surfaces, which eliminates the previous design restrictions on implant components.The articulating surfaces of a knee implant have a generic shape that might change the gait for many patients.

These have been well-known problems for a long time, and in the ‘70s, several attempts were made to develop custom-sized implants that would fit better. However, the fabrication technologies available at that time made it prohibitively expensive to produce custom parts. The custom sized implants were based on radiographs, which was a fairly inaccurate method due to the 2-dimensional limitations. Most modern orthopedic implants are manufactured through investment casting, which is an excellent fabrication technology that provides high precision parts with good surface finish. Unfortunately it is expensive in small quantities.

Unfortunately it is expensive in small quantities. Solid freeform fabrication (SFF) technologies have enabled manufacturers to produce small quantities through investment casting at a reasonable price, opening up new possibilities for custom design and fabrication of orthopedic implants.Recent advancements in medical imaging and image processing have enabled custom design of biomedical implants based on patient specific Computed Tomography (CT) data [11,12]. The CT data is edited, and an accurate 3D-model of the joint is created and exported as an stl-file. A different software is used to convert the stl-file into a CAD-model that can be used as the base for the custom designed implant components. Several SFF-technologies can be used to produce a master pattern for investment casting, but producing a finished implant component is still time consuming and labor intensive.

Several SFF-technologies can be used to produce a master pattern for investment casting, but producing a finished implant component is still time consuming and labor intensive. With the introduction of the Electron Beam Melting (EBM) machine by Arcam (Sweden), a new possibility for fabrication of custom implant components has become available. Initially, the EBM technology was only available with tool steel, which is not a biocompatible material. In theory, the EBM machine can process most440materials that are electrically conductive, and a collaborative effort between North Carolina State University and Arcam AB was initiated to develop titanium for the EBM process.To enable clinical testing of the custom designed implants, an animal study is the first step, and collaboration between the Industrial Engineering Department and the Veterinary School at NCSU was initiated [13].

Surgeons at the veterinary school have developed generic hip implants for canines that are currently being implanted on a weekly basis. At this time, no commercial knee implants for canines are available due to the number of sizes that would be needed to accommodate most patients. However, the demand is present. The idea of custom designing and fabricating knee implant components for canines became an interesting solution for the veterinary surgeons at the same time as it would serve as a clinical trial for humans. A CT-scan of a potential patient was acquired, and a custom femoral component was designed in collaboration with the orthopedic surgeons. Both the articulating surface and the bone-implant interface surface were custom designed based on the CT-data.This paper describes the development of the titanium powder for the EBM process as well as the initial fabrication of the first custom implant components.

The traditional investment casting process of implant components is compared to the EBM fabrication.Figure 2. Computer models of a custom designed implant component for a canine stifle joint.Material and MethodsAs a first step to enable direct fabrication of custom designed orthopedic implants using the Electron Beam Melting technology, the process parameters for titanium (Ti6Al4V) were developed. North Carolina State University partnered with Arcam AB in the development, which took place at Arcam’s facility in Mölndal, Sweden. Both Ti6Al4V and pure titanium are commonly used for biomedical implants. Pure Ti is softer and less suitable for high impact applications. The first challenge was to find a supplier that offers a Ti powder with the right composition and particle size, which cannot be disclosed at this time. The particle size is very important for the final result and affects the flow of the powder and the melt pool behavior.

The particle size is very important for the final result and affects the flow of the powder and the melt pool behavior. For each new material, a long list of processing parameters must be developed through experimentation. The initial parameters are calculated based on material dependent factors such as melting temperature and flow characteristics. Each parameter is optimized through an iterative process based on testing and evaluation. To further complicate matters, the process parameters are geometry dependent as well and require a vast amount of testing to fully develop. At the441present time, parameters for many types of geometries have been developed and successfully fabricated, however no custom designed implant component has been fabricated in titanium. The processing parameters and the powder composition for the Ti6Al4V is proprietary and cannot be disclosed at this moment but the processed material properties are reported in the next section.Figure 3.

Chess piece built of Ti6Al4V in the EBM S12The Electron Beam Melting process is powder based, and the excess powder supports the structure during the build phase. It does not eliminate the need for support structure, however. The parts are built on a substrate, and the part and start plate are kept at an elevated temperature throughout the build to minimize residual stresses. When the build is completed, the part and substrate will cool down and a shrinkage of 1.4 % for tool steel will take place that has been compensated for through the software prior to the build. The part tends to shrink towards the center causing a curling or bending motion of the edges. To prevent distortion, support structure is used to secure the part to the substrate and to secure overhanging structures to the underlying structure. The orientation of a part is very important considering the amount of support structure needed and the surface finish of the final part.

The orientation of a part is very important considering the amount of support structure needed and the surface finish of the final part. On thin features with down facing surfaces, “icicles” can form that will affect the surface finish and complicate the finishing of the part. At the present time, no automatic support generation for the EBM process is available. If Magics from Materialize is used to generate support, substantial editing is needed since excessive support is generated. For this project, the implant component was oriented on its side, and support structure was manually created directly in SolidWorks. When building parts with very thin cross sections, it is difficult to effectively transmit enough heat through the part to keep the elevated temperature without over melting the powder. To solve this problem, a “dummy” structure was added as an artificial means to impart more heat to each.

To solve this problem, a “dummy” structure was added as an artificial means to impart more heat to each. In this case, a large diameter cylinder was added to increase the total melting area for each layer.Figure 4. Computer model showing the implant with support and heat sink442As a time and cost comparison, a canine stifle joint component was fabricated using the Electron Beam Melting machine and was compared with traditional investment casting using an RP-pattern as a master. It was decided to use an SLA-190 to fabricate the master pattern to be used for the investment casting process and a QuickCast model of the implant was produced. Due to the small and thin geometry of the canine femoral component, it was difficult to drain the resin from the model leaving an almost solid RP-pattern. Several attempts of reorienting the component and creating drain and vent holes still did not deliver a hollow structure sufficient for the QuickCast process.

Several attempts of reorienting the component and creating drain and vent holes still did not deliver a hollow structure sufficient for the QuickCast process. To complete the investment casting process, an RTV-mold was created using the SLA-pattern and a wax pattern was cast.Figure 5. RTV mold with SLA RP-master (L) and resulting wax pattern (R)The wax pattern was attached to a wax sprue and placed in an investment-casting flask. Quikvest was used as the ceramic media and poured around the wax pattern. The flask was placed in a convection oven to remove moisture and later fired in a high temperature furnace according to the manufacturer’s recommendation. For this comparison, aluminum was used for the casting.ResultsThe material properties of the EBM processed titanium powder are very similar to conventionally processed titanium. The processed titanium is fully dense and is biocompatible in accordance with FDA regulations.

The processed titanium is fully dense and is biocompatible in accordance with FDA regulations. The Ti6Al4V parts fabricated in the EBM process feature good machinability and can be machined as stock Ti6Al4V parts. Pure titanium is under development but the material properties are not publishable at this time. The chemical composition of the EBM processed Ti6Al4V is as follows:Element Nitrogen (N) Oxygen (O) Aluminum (Al) Vanadium (V) Iron (Fe) Nickel (Ni) Chromium (Cr) Sulphur (S) Manganese (Mn) Tin (Sn) Zirconium (Zr)Weight- % 0.01 0.19 6.2 4.5 0.45 0.037 0.03 < 0.01 0.04 1.3 < 0.02Table 1. Typical chemical composition of EBM processed Ti6Al4V (Source: Arcam Ti6Al4V Data Sheet)443The mechanical properties of the EBM processed Ti6Al4V are as follows:Property Hardness HRc Tensile Strength, RmYield Strength, Rp0.2Elongation30-35 HRc 930 MPa 135 000 psi 880 MPa 120 000 psi 10%Table 2.

Mechanical properties of Ti6Al4V (Source: Arcam Ti6Al4V Data Sheet)At the time of writing this paper, the processing parameters for the Ti6Al4V were not fully developed for thin walled structures like the stifle joint component. As a proof of concept, the custom designed stifle joint implant was fabricated on the EBM machine using the commercially available tool steel H13. The total building time for the implant was 5 hours. After the part had cooled down to room temperature, the support structure was removed using a rotary cutting tool and the articulating surface was polished to a mirror finish. The total finishing time was approximately 2 hours, which would be similar if it had been fabricated out of titanium. The total setup time for the EBM machine was approximately 2 hours and the implant was left to cool down over night.Figure 6. Custom designed knee implant processed on EBM S12The investment-cast implant took considerably longer time to complete with many more steps.

Custom designed knee implant processed on EBM S12The investment-cast implant took considerably longer time to complete with many more steps. The total time for the investment cast implant from stl-file to finished product was 78 hours compared to 25 hours for the EBM process. A considerable amount of time was spent waiting for the RTV mold to cure as well as drying the investment. A total time of 11 hours of labor was spent on the investment cast version compared to 7.25 hours for the EBM version. The total machine time for the EBM machine is 8.25 hours as compared to 8.5 for the SLA machine.

The hourly rate for an EBM machine is slightly higher than a modern SLA machine, but the investment cast version requires additional equipment such as a convection oven, a furnace for firing the investment, as well as melting the metal, and additional tools needed for the mixing of investment and RTV rubber.444Investment Casting Start with STL Prepare SLA Build File Startup SLA SLA build time Clean and postcure Sand SLA pattern Prepare to Pour Silicone Silicone Cure time Finish silicone mold Wax casting Build wax tree Mix and pour investment Dry investment Fire investment Pour metal Cool down Cut off tree Clean/Polish TotalHours EBMStart with STL1 Prepare EBM Build File 0.5 Startup EBM Machine 8 Draw Vacuum 3.5 Heat Start Plate 0.5 EBM Build Time 1 Cool down 24 Remove sintered powder 0.5 1 0.5 0.5 24 10 0.25 10.25 Cut supports/ remove start plate 1.5 Clean/Polish 78 TotalHours1.5 2 0.75 0.5 5 12 0.251 2 25Table 3.

Time comparison between investment casting and EBM processing of custom designed implantDiscussion and ConclusionsIt is the authors’ opinion that the EBM technology can successfully be used to fabricate custom designed implants for knees, hips, elbows, shoulders, fingers, and bone plates in titanium. Even though the initial results appear promising, additional material development will be necessary to optimize the processing parameters for Ti6Al4V and Ti. There are differences between processing steel and titanium using the EBM technology. The build time for titanium is approximately half of that for steel, however the cool down must take place under vacuum inside the build chamber. There are other potential advantages with using the EBM technology to fabricate custom designed orthopedic implants that can not be achieved through traditional processes such as investment casting.

There are other potential advantages with using the EBM technology to fabricate custom designed orthopedic implants that can not be achieved through traditional processes such as investment casting. In the case of a knee implant component, it is desired to have a very hard and smooth surface finish on all articulating surfaces and a soft and porous bone-implant interface. To achieve this through investment casting, a porous coating is applied to the bone-implant interface surface through a sintering process. Similar coatings are applied to hip, elbow, and shoulder implants to promote the bone ingrowth in selected areas. The Electron Beam Melting technology has the prospect of fabricating parts with functional gradient microstructures, which would be very useful for orthopedic implants. This would allow for the porous surfaces to be directly designed into the components.

This would allow for the porous surfaces to be directly designed into the components. Further, the EBM technology lends itself to deposit multiple materials that is of highest interest to the orthopedic implant industry as well. Often orthopedic implant components are made out of several materials to achieve the desired properties. Modular hip implant systems are often combined out of both titanium and cobalt-chromium components.445References1. Taylor J, Rorabeck C H, Bourne R B , et al: Total Knee Arthroplasty in Patients 50 Years or Younger: Long-Term Follow-Up. American Academy of Orthopedic Surgeons 2000, Paper No 185, Orlando, AAOS.2. Knutson, K., S. Lewold, et al. (1994). "The Swedish knee arthroplasty register. A nation- wide study of 30,003 knees 1976-1992." Acta Orthop Scand 65(4): 375-86.3. Robertsson, O., K. Knutson, et al. (1999). Knee Arthroplasty for Osteoarthrosis and Reumatoid Arthritis 1986-1995. 1999 American Academy of Orthopaedic Surgeons, Anaheim.4.

rect Fabrication of Polymer Composite Structures with Cuwed LOMDonald Klosterman, Richard Chartoff, Mukesh Agarwala Ira Fiscus, John Murphy, Sean Cullen, Mark YeazellRapid Prototype Development Laboratory University of Dayton 300 College Park Dayton, OH 45469-0 13 1ABSTRACTThis report describes the application of Curved LOM to the direct fabrication of polymer matrix composites (PMCs). The overall methodology of directly fabricating PMC parts involved the use of the Curved LOM machine to lay-up and shape "green" composite laminates from prepreg feedstocks, followed by vacuum bag / oven cure and consolidation. The conventional Curved LOM laminator was replaced with a vacuum thermoforming apparatus to better accommodate the bonding of commercially available prepregs. The study also demonstrated that it is possible to interface a general composite design sofiware package with the Curved LOM machine via the curved slice file (.CSF) format.

The study also demonstrated that it is possible to interface a general composite design sofiware package with the Curved LOM machine via the curved slice file (.CSF) format. Taken together, these two improvements allow for improved flexibility in manufacturing PMC components, from both a material handling and a design point-of-view. A simple C-shaped panel was fabricated and tested to demonstrate the overall feasibility of the process for PMCs. A glass fiber 1 epoxy prepreg obtained from a commercial supplier was used as a model material system.

A glass fiber 1 epoxy prepreg obtained from a commercial supplier was used as a model material system. It was found that the cumulative accuracy of the overall process was good, and the mechanical properties of the laminates were acceptable for non- structural applications for which the material is normally used.INTRODUCTIONThe Curved-Layer Laminated Object Manufacturing process (Curved LOM) was developed fordirect, automated fabrication of curved layer structures made from high temperature materials such as ceramic matrix composites (CMCs) or monolithic ceramics. Full description of the process can be found in previous publications [1,2,3]. Described in this report is the application of Curved LOM to the direct fabrication of polymer matrix composites (PMCs). The availability of a direct method for rapidly and inexpensively producing PMC prototype components would be of great benefit to the composites design and manufacturing community.

The availability of a direct method for rapidly and inexpensively producing PMC prototype components would be of great benefit to the composites design and manufacturing community. Currently, the effort and cost involved in the design and fabrication of PMC prototype components can be great. The design community needs a rapid, low-cost prototype manufacturing capability that can fabricate prototype parts of comparable quality to those produced via traditional techniques. In fact, aerospace designers have been recently turning to rapid prototyping technologies for potential new solutions [4]. Using LOM for the lay-up of PMC prototype laminates will allow for the simultaneous evaluation of new materials and design concepts, enabling rapid design and fabrication of competing prototypes.The purpose of this report is to describe recent improvements made to the hardware and software capabilities of the Curved LOM apparatus (see [3] for a previous status report).

These updates were made to allow for improved flexibility in manufacturing PMC components, from both a material handling and a design point-of-view. A simple C-shaped fiber glass I epoxy panel was fabricated to demonstrate the overall feasibility of the process for PMCs. Using this demonstration component and other fabricated flat panels, the dimensional accuracy of the overall process was determined, and the quality of the panels was assessed through measurements of the fiber volume fraction and interlaminar shear strength.Improvements were made to the Curved LOM apparatus to better accommodate commonlyavailable PMC materials. Preimpregnated fiber preforms, or "prepregs", are widely used by manufacturers of composite products. Prepregs are fiat sheets containing fibers (usually continuous unidirectional or woven fabrics) and a polymer resin. It is common industrial practice to build parts by cutting the prepregs to shape and laying them up manually.

It is common industrial practice to build parts by cutting the prepregs to shape and laying them up manually. Layers are compacted manually using a roller and/or vacuum diaphragm in order to dispel air from in between layers. This compaction capability was not available on the initial version of the Curved LOM machine [3]. Thus, a vacuum thermofoming apparatus was developed in order to improve on total pressure available and pressure uniformity.The new laminator design is illustrated in Figure 1. A flexible, resistively heated pad is sandwiched between two silicone rubber diaphragms. The composite diaphragm is stretched over and secured to the open side of a 13 inch x 13 inch x 3 inch aluminum box. This mechanism is shuttled directly over the LOM-paper mandrel, which is supported by a vacuum plate. A prepreg is placed over the mandrel, and the build platform is elevated until it seals with the laminator.

A prepreg is placed over the mandrel, and the build platform is elevated until it seals with the laminator. A vacuum is drawn through a perimeter channel that has been bored in the vacuum plate. A porous "breather ply" (not shown), such as a woven fiberglass layer, can be placed over the lay-up to ensure even vacuum distribution over the entire part area. With a breather ply in place, the vacuum pressures at the top center and far edge of the part were measured to be 1 1.8 psi and 12.0 psi, respectively. It is possible to provide additional pressure via compressed gas using the vent tube. Currently, the maximum size part that can be made is 8 inch x 8 inch x 3 inch. Given the envelope of the LOM machine housing, it is possible to scale- up to 30 inch x 20 inch x 5 inch parts.The interaction of sofhvare programs necessary for Curved LOM is illustrated in Figure 2 (see [l]for an expanded description). The algorithms are incorporated in two separate packages.

The algorithms are incorporated in two separate packages. First, LOMSlice 2.0 is used to slice an .STL file into a series of curved layers. The information of each layer, namely the X Y Z coordinates of the layer cross section, is stored in a separate .CSF (Curved Slice File) file. Thus, each part is represented by a series of .CSF files, equal to the total number of layers required to build the part. Next, the .CSF files are exported to 3D LOMSlice which drives the Curved LOM machine via an eight-axis motion control card.This scheme is problematic for building PMC parts, based on the inherent shortcomings of the .STL file format. Because the .STL format is a surface representation, it cannot contain layer-by-layer information such as ply architecture (i.e. fiber direction) needed to fully describe a composite structure. Furthermore, engineers who design composite structures are typically not familiar with the .STL format.

fiber direction) needed to fully describe a composite structure. Furthermore, engineers who design composite structures are typically not familiar with the .STL format. On the other hand, the .CSF format is of the appropriate geometrical format, although it is also relatively unknown to composite designers. One purpose of this study was to demonstrate the potential use of .CSF files as a "neutral exchange" format, providing the link between the Curved LOM machine and softwarelfile formats typically used by composite designers.The design of composite structures is normally accomplished with suites of software packagesthat have been developed to accommodate all aspects of composite design, such as ply shape definition, prepreg orientation and stacking sequence, and structural analysis and optimization [e.g. 51.

51. Through simulation, designers determine the best ply architecture (i.e., shape, sequence, and orientation of prepreg layers) for a given part subjected to the expected loading conditions. Some software packages [6] have the additional capability of providing a direct link from 3D CAD geometry to automated equipment, such as prepreg cutters and laser projection systems used on the manufacturing floor.Based on this existing capability for direct machine control, it is not difficult to envision the possibility of controlling the Curved LOM machine by these packages. The motions of the various Curved LOM hardware elements are controlled by an off-the-shelf machine control board (Galil DMC- 1000). However, instead of having commercial design software packages interface directly with this DMC card, it may be easier to output .CSF files. Illustrated in Figure 3 is an example of a software package that accomplishes this task.

Illustrated in Figure 3 is an example of a software package that accomplishes this task. This package, referred to as Curved Composite Panel Designer (CCPD), creates the ply information for simple "C" shell composite panels. The user inputs the desired width, height, length, shell thickness, layer thickness, and layer-by-layer ply orientation of a simple "C" panel containing parabolic surfaces. The program outputs a series of .CSF files, each containing the XYZ coordinates and ply orientation of a single layer. The program is also capable of creating a simple "tool outline" file, which greatly simplifies the accurate placement and registration of the LOM-paper mandrel in the Curved LOM machine. In a similar manner, it is thought that commercially available software design packages could be readily augmented with a "Curved LOM" output module or "translator" to perform the same function, as illustrated in Figure 2.

Since such software programs are CAD-based, it would also be possible to output the necessary .STL files for creating the curved mandrels, which can then be built with flat LOM or other SFF techniques. Overall, the benefit of such an effort would be to drastically improve the availability of rapid prototyping methods such as LOM to the composite design community.PMG PIIOTOTWE FAEMCATIONThe methodology for direct fabrication PMC components with Curved LOM is illustrated inFigure 4. Prepregs are used as feedstocks due to their wide commercial availability and ease of processing. The LOM machine is used as a means of automated lay-up and cutting, ideally eIiminating touch labor and improving accuracy and consistency. Throughout the process, some heat can be used to assist lamination if necessary. However, it is desirable to keep the prepregs and the part in a state of minimal cure.

Throughout the process, some heat can be used to assist lamination if necessary. However, it is desirable to keep the prepregs and the part in a state of minimal cure. It is advantageous to perform most of the cure and consolidation in one step afler the part is completely layed-up. This approach will ensure that layer compaction, resin flow (which assists compaction), and the majority of the resin polymerization will occur simultaneously and uniformly throughout the part, as a result of the steady pressure and temperature conditions. This consolidation cycle can be performed using a variety of techniques, such as an autoclave cure or matched die press molding, but vacuum bag / oven cure was chosen to remain consistent with the idea of low cost processing.A simple "C" shell panel was fabricated for the purpose of demonstrating the overall CurvedLOM fabrication of a PMC, including the new hardware and software capabilities.

A glass fiber / epoxy prepreg, obtained from a commercial supplier [7], was used as a model PMC material system. The prepreg was comprised of a woven E-glass fiber mat (plain weave, style 7628) impregnated with a B- staged (partially cured), catalyzed Novolac epoxy resin system. The resin content and thickness of the prepreg were 35 wt% and 0.0085 inches, respectively. The prepreg was slightly tacky to the touch at room temperature, and thus it was supplied with a removable backing ply (thin plastic sheet, like cellophane).First, a LOM-paper mandrel was created to support the "C" shell. An .STL file for the LOM-paper mandrel was created using the .STL file of a "C"shel1 part and LOMSlice 2.0 software, as illustrated in Figure 2. The mandrel was built on a LOM 101 5 machine using Helisys LPH008 paper.

The mandrel was built on a LOM 101 5 machine using Helisys LPH008 paper. Next, the dimensions of the "C" shell, the average cured thickness of the prepreg (0.007 inches), and the lay-up sequence (16 ply, quasi-isotropic), were input to CCPD software program. The resulting sixteen .CSF files were loaded into 3D LOMSlice. The LOM-paper mandrel was placed on the Curved LOM build platform. The prepregs were layed-up on the mandrel by hand (the automatic sheet feeder was disassembled due to redesign and implementation of the new laminator). A mylar film and woven glassfiber mat (Crowfoot weave, Style 120) were placed over each layer to serve as release and breather plies, respectively. Each layer was laminated for 30 seconds at room temperature. After lamination, the mylar and breather plies were removed prior to laser cutting.

Each layer was laminated for 30 seconds at room temperature. After lamination, the mylar and breather plies were removed prior to laser cutting. All sixteen plies were processed without problem.The finished green part with attached LOM-paper mandrel were removed in one piece from the Curved LOM machine, placed on an aluminum sheet, vacuum bagged [8], and placed in a temperature- controlled oven. The cure cycle was 25 to 1 5Q°C in one hour, three-hour hold at 1 5Q°C, and free cool to 25OC. The vacuum pressure was maintained at >14 psi for the duration. The resulting part and a similarly fabricated body armor panel are shown in Figure 5. During the cure cycle the LOM paper mandrel shrunk by 11% in the z direction, and zero in the width and length. The "C9' shell experienced spring-back (which is normally encountered in curved PMC panels), and thus, its dimensional change was much less than the mai~drel.

The "C9' shell experienced spring-back (which is normally encountered in curved PMC panels), and thus, its dimensional change was much less than the mai~drel. In fact, the cured "C" shell fit almost perfectly on a freshly-built mandrel, as illustrated in Figure 5.DIMENSIONAL ACCURACY AND MECHANICAL PROPERTIESThe cumulative dimensional accuracy of the process is summarized in Table 1. No attempt wasmade to incorporate shrinkage compensation factors into the original dimensions. All measured final dimensions except height were within 1% of the design specification. Two possible sources of error for the height dimension are: LOM process-related error (likely due to platform motion and level control) and cure shrinkage in the direction of curvature. Due to the reinforcing fibers, it is unlikely that the second consideration is significant. Tl~us, additional study of the platform motion and control is needed.

Due to the reinforcing fibers, it is unlikely that the second consideration is significant. Tl~us, additional study of the platform motion and control is needed. On the other hand, the laser gantry X Y motions seem to be highly accurate, as manifested by the small deviations in the length and width dimensions.Table 1: Cumulative dimensional accuracy of glass fiber / epoxy "C9'shell part made with Curved LOM / vacuum bag oven cure process.Edge burning can clearly be seen on the PMC parts, due to charring from the cutting action of the C 0 2 laser. Figure 6 shows a close-up of a laser cut line on the glass fiber / epoxy prepreg. The total laser kerf, including char zone, is about 0.5 mm. Previous studies have shown that this degradation can be minimized or eliminated only by using other types of lasers capable of photoablation [9].

Previous studies have shown that this degradation can be minimized or eliminated only by using other types of lasers capable of photoablation [9]. The effect of the charred edges on the mechanical properties of the resulting panels is unknown.I11 general, the fiber volume fraction in composite panels serves as a primary point of comparison.Based on the cured thickness, number of plies, prepreg fiber areal weight, and fiber theoretical density, the fiber volume fraction was calculated to be 44%. Given the method of consolidation, i.e. vacuum bag / oven cure, this result was promising. Furthermore, this number was moderately repeatable (4 1-45 volume%) for several flat and curved panels made from the same prepreg in a variety of sizes (2.5"x4" to 6"x6", 5 to 18 layers thick). In high performance aerospace applications, coinposites with fiber volume fraction of 60-70% are required. However, this usually necessitates the use of expensive autoclaves for consolidation and cure.

However, this usually necessitates the use of expensive autoclaves for consolidation and cure. Thus, the LOM 1 vacuum bag processing route seems to be a feasible method for quick, low-cost manufacture of reasonable quality PMC parts. If higher fiber volume fraction is desired, the mcured LOM laininates could be consolidated in an autoclave.The interlaminar shear strength test [lo] is widely used as a screening tool for PMCs. Flat panelsof two different thicknesses were fabricated on the Curved LOM machine using the glass fiber 1 epoxy prepreg, followed by the vacuum bag I oven cure cycle as described above. Test coupons were cut from the cured panels using a diamond cutter. The test results are given in Table 2. It is often difficult to size the specimens such that a delaminationlshear failure is observed in the test. In the present case, two different sized specimens failed in the proper mode, with almost identical match in the calculated interlaminar shear strength.

In the present case, two different sized specimens failed in the proper mode, with almost identical match in the calculated interlaminar shear strength. Another specimen failed at a higher strength, but the observed failure mode was not shear. Thus, it would be prudent to report a value of approximately 3600 psi as the measured shear strength.Table 2: shear strength results' for flat panel, glass fiber 1 epoxy composites fabricated with Curved LOM and vacuum bag 1 oven cure cycle.diameter loading pins, and room temperature.No manufacturer data was available on the shear strength properties of this prepreg. In general the highest performance composites, fabricated from structural resins and consolidated with autoclaves, have shear strength values as high as 12,000-1 8,000 psi.

In general the highest performance composites, fabricated from structural resins and consolidated with autoclaves, have shear strength values as high as 12,000-1 8,000 psi. Possible reasons for the shear strength of the current material being lower than this are the following: the resin is inherently not suited for structural applications; vacuum bag consolidation does not lead to the highest possible fiber volume fraction; and the span:depth ratio (8: 1) was larger than normally used in the short beam shear test (5: I), which will reduce the apparent shear strength. However, the normal use of the current material is in abrasion resistant, electrical applications, such as bearing insulation.

However, the normal use of the current material is in abrasion resistant, electrical applications, such as bearing insulation. It is thought that the shear strength is adequate for such applications.RELATED mSEARCI-TSignificant related research has recently been completed that provides science-based tools toquantify the ability to laininate PMC prepregs on the Curved LOM machine, predict layer compaction, and predict the extent of cure during LOM processing. These results are the subject of future publications [I 1, 12, 131. This new approach will eliminate trial-and-error methods of determining the suitability of new materials in the Curved LOM process, as well as guide the user in determining machine operating conditions.CONCLUSIONSHardware and software for the Curved LOM Process was upgraded to improve the ability fordirect fabrication of polymer composite prototypes.

A simple "CC" shell was fabricated from a commercially available prepreg to demonstrate the new capabilities. The Curved LOM process was used to fabricate the "green" (uncured) part, and a vacuum bag 1 oven cure cycle was used to provide final consolidation and cure. The accuracy of the overall process was found to be good, as most dimensionswere within 1% of the design specification without incorporating shrinkage compensation factors. The interlaminar shear properties of the cured laminates was judged as acceptable for the normal applications of the material.The authors wish to thank Dr. William Coblenz (DARPA) and Dr. Steven Fishman (ONR) whohave supported the development of Curved LOM over the past several years. Application of Curved LOM to polymer matrix composites was supported by a Research Challenge Grant from the Ohio Board of Regents. The authors would like to thank Mr. Jerry Fleeson of Lectromat, Inc. for donating the polymer pregpreg used throughout the study.

The authors would like to thank Mr. Jerry Fleeson of Lectromat, Inc. for donating the polymer pregpreg used throughout the study. The authors would also like to acknowledge Mr. Don Byrge for mechanical testing advice and Mr. Gary Andrews for specimen preparation.m F E m N C E S1. Klosterman, D., R. Chartoff, N. Osborne, G. Graves, A. Lightman, G. Han, A. Bezeredi, S. Rodrigues, "Development of a Curved Layer LOM Process for Monolithic Cermics and Ceramic Matrix Composites," Rapid Prototyping Journal, Vol. 5, Issue 2, 1999.2. Klosterman, D., R. Chartoff, N. Osborne, G. Graves, A. Lightman, G, Han, A. Bezeredi, S. Rodrigues, "Development of a Curved Layer Process for Fiber-Reinforced Composite Materials," Eighth European Conference on Rapid Prototyping and Manufacturing, Nottingham, UK, July, 1999.3. Klosterman, D., R. Chartoff, N. Osborne, G. Graves, A. Lightman, G. Han, A. Bezeredi, S.

8. According to Standard Guide for Preparation of Flat Composite Panels with Processing Guidelines for Specimen Preparation, ASTM Standard D5687 / D5 687M-95.9. Lightman, A., G. Han, "Laser Cutting of Ceramic Composite Layers," ," Solid Freeform Fabrication Symposium Proceedings, University of Texas at Austin, Austin, TX, August, 1996, pp. 291-298. 10. Standard Test Method for Apparent Interlaminar Shear Strength of Parallel Fiber Composites by Short-Beam Method, ASTM Standard D2344 - 84.1 1. Popp, A., D. Klosterman, R. Chartoff, "Development of Science-Based Tools for Material and Process Considerations in Curved LOM, Part I: Application and Use of a Tack Test," Rapid Prototyping Journal, manuscript in development.12. Bryant, E., D. Klosterman, L. Flach, R. Chartoff, "Development of Science-Based Tools for Material and Process Considerations in Curved LOM, Part 11: Development and Verification of a Thermal / Cure Model" Rapid Prototyping Journal, manuscript in development. 1 3.

The use of CSF files allows for direct use of general/commercial composite design software packages, by-passing the STL file.Figure 3: example of general composite design package, Curved Composite Panel Designer (CCPD), in which composite ply information is output as XYZ coordinates.Green lamin;anteFigure 4: Curved LOM processing methodology for polymer composites, involving use of Curved LOM to form "green" (uncured) panels which are subsequeiitly cured in a low-cost vacuum bag 1 oven cure cycle.IFigure 5: (above) cured, glass fiberlepoxy composite parts made with Curved LOM, and (below) cured, curved layer part resting on fresh LOM-paper mandrel.0.5 rnrn kerf (including burn-back zone)Figure 6: photomicrograph of glass fiber 1 epoxy prepreg cut with C02 laser on Curved LOM machine..

rect Feedback Control of Gas-Phase Laser-Induced DepositionJ. l. MaxweU1, J. Pegna, D. Messia, D. DeAngetis Center for Integrated Electronics and Integrated Manufacturing, Rensselaer Polytechnic Institute, Troy, NewYork 12' 80·3590Tel: (518) 276-6030, Fax: (518) 276-2623, email: pegnaj@rpLedu.Abstract: Three-dimensional laser Chemical vapor deposition (3D-LCVD) or SALD, was used to prototype metallic and ceramic microstructures. Iron, nickel, and steel metal forms were grown from organic and halogen based precursors. Through the simultaneous use of multiple precur sors, specific nickel-iron based alloys were produced. By observing the emission spectra during growth, a measure ofthe volumetric growth rate, was obtained. Direct, PID control ofthe process was then possible using the growth rate measurement as real-time feedback.

Direct, PID control ofthe process was then possible using the growth rate measurement as real-time feedback. Calibrated infrared photographs of evolving microstructures were taken at various wavelengths, giving a measure of the temperature gradient over the growth zone. While radiation contributes to heat losses at high temperatures, enhanced convection is the dominant heat transfer mechanism due to the small dimensions of the heated area. Enhanced growth rates, induced by convective flow, were also observed. The heat and mass transfer coefficients were determinedfor various processing condi tions, and compare well with experimental data.

The heat and mass transfer coefficients were determinedfor various processing condi tions, and compare well with experimental data. Axi-symmetric rods may also be grown in both the kinetic and transport-limited regimes, and a systematic study of the precursor pressure and deposit temperature during growth yielded distinct growth regimes, influn, ced by the interplay of heat losses and diffusive transport.Keywords: Solid Area Laser Deposition, Chemical Vapor Deposition, Grc..'phite, Iron, Nickel, Nickel-Iron Alloys, PID Process Control.1. Current Affiliation: Institute for Micromanufacturing, Louisiana Tech University, Ruston, La 712701 INTRODUCTION Laser-induced chemical vapor deposition (LCYD) has· been explored extensively as a thin-film metallization tool. In contrast, three-dimensional laser chemical vapor deposition, 3D-LCYD (or SALD), is a mode of pyro lytic deposition where more elevated structures can be generated, as exemplified in figures 1, 2, and 3.

Our interest in 3D-LCYD is in its use as a free-form rapid-~~r::::~:t:~:::::rc~~:::f~:~l~~~::~C;:::::..have been for the most part open-loop, growth rates and ii:ii~::::!:i:il;:f::::::::: required process conditions being estimated from empir-;;;:::;:::j':::ll;:n:::::\~?: ical data. The purpose of this work was to determine ali::!::}}:;;;;':::::;':':':::':':':': method for measuring and controlling the volumetric::::~: C7 deposition rate during 3D-LCYD. =FI:-:G U:-::R:-::E:-:1-.----Sample Nickel Structure.<I:> 1996 Moxwe1l, Pcgll4. M",i. and DeAngelis printt<l with pamissioo.227In this work, the volumetric growth rate was measured directly from the emission spectra asso ciated with the pyro lytic reaction. This rate measurement was used to control the input power and instantaneous growth rates. Applications include control of laser direct writing over varying su b strate topography, control of undesirable ==== features in layered FIGURE 2..

Applications include control of laser direct writing over varying su b strate topography, control of undesirable ==== features in layered FIGURE 2. . growth, and Imple- mentation of continuous 3-dimensional deposition.structure.~~~~gl;i[~~ji~t~~:i1ff100¥f~m!···;···:···:···1F 33 mharc 86 mhar~~~~:~~u~~~~~~~~:~~i~:~O~ee~~~~:i~~~)~~ the iron and nickel carbonyl precursors. This result, ~ however, will not be discussed within the scope of this paper. For more information on this subject, the reader is referred to [Maxwell et aI., 1996].~ 10~~~::M~:·:':'~:·::·~·;··;;f::::~·1":::i:~:·f::::~4::~~1o.4008001'2001600H 132 mharF 168 mharIncident Laser Pnwer*2 EXPERIMENTAL The 3D-LCVD reactor at Rensselaer consists of a custom quartz tube with a port for viewing and laser input. The chamber is connected to a pumping station via a gate valve.

The chamber is connected to a pumping station via a gate valve. The vacuum chamber and gas-delivery system are enclosed within a venti lated hood for safety purposes.J:::iJ;I"~~1 J :::·;T+HH1iilJilF 168 mharc 132 mharH 33 mharIron pentacarbonyl was delivered to the reac tion chamber via an evaporator. In this experiment, the room-temperature vapor pressure of Fe(CO)S was used to fill the chamber. In most cases, the par tial pressure of the precursor reached 20 mbar.o40080012001600Incident Laser Power*Corrected for optical losses.- _.. _._---:--;-----;---:- Nickel average axial growth ratesFIGURE 4. and diameter vs. input powt3r.228Both ethylene and nickel tetracarbonyl were metered into the reactor from commercial cylin ders. For the growth of pyrolytic graphite, ethylene pressures of up to 930 mbar were employed. Ni(CO)4 partial pressures of up to 250 mbar were possible. Iron pentacarbonyl was delivered to the reaction chamber via an evaporator.

Ni(CO)4 partial pressures of up to 250 mbar were possible. Iron pentacarbonyl was delivered to the reaction chamber via an evaporator. In these experiments, the room-temperature vapor pres sure of Fe(CO)S was used to fill the chamber, i.e. 20 mbar.The beam source was a Spectra Physics argon-ion laser, with a maximum output of 8 watts (multi-mode) at the 488/514 nm primary lines. AtJ Diameter[pm]these wavelengths, iron and a:::l. ........ nickel have normal spectral reflectances of 65% and 52% .§ respectively [Touloukian, 1970]. Gaussian spot radii of 25 ~m (l/e2) or greater were used throughout the experi ments, and it was found that while the commencement of growth on the substrate FIGURES. depended greatly on the power density-requiring small spot sizes for the laser powers available-- the 3-D, steady-state growth of rods was nearly unaffected by the beam waist over a distance several times the Rayleigh range (250 ~m).

The optical system consisted of a Newport lOx beam expander and 200 mm focal length CVI achromat. A long pass filter was used to block the UV laser lines below 420 nm.o Length [pm]tI)1::]~~~~~!'~~l~~T~'~I~~~I~'~~f~'~~~F!f~:~r1Jj~~ :~'~~f~~:~~!~~~~:~:1~~o 1000 500 Time from Initiation of Growth [s]250750- - -Nickel rod length and mid-length diameter vs. timeA silicon photodetector, covered with narrowband filters, was mounted to one of the chamber windows, at a distance roughly 200 mm from the sample. Using a two-d~cade pre-amplifier, the sensor could measure emissions as small as 0.25 mW at the substrate (at 656nm). In this arrange ment, the sensor had a time response on the order of milliseconds; however, the amplifier was heavily filtered with a time-constant of approximately 1/30 s.

In this arrange ment, the sensor had a time response on the order of milliseconds; however, the amplifier was heavily filtered with a time-constant of approximately 1/30 s. The amplified signal was recorded by an Omega Nubus data acquisition system, with a typical sample period of 0.05-0.10 seconds, and was later time averaged as needed for real-time control. Proportional-Integral-Derivative (PID) Control of the laser power was implemented using Omega Workbench™ software with out put to the LCD retarder mentioned previously.For analysis of the sample composition, both Auger electron spectroscopy (PHI 545C Scan ning Auger Microprobe) and an electron microprobe (Geol Superprobe 733) were employed.

Prior to the microprobe analysis, the samples were sectioned, polished, and lightly etched in a Kalling's solution, a mixture of isopropyl alcohol, HCL, and CuCI2.2293 RESULTS AND DISCUSSION3.1 NI GROWTHNi(CO)4 vapor pressures ranged from 25 to 200 mbar throughout the experiment, and rods could be initiated at pressures of less than 25 mbar. At similar partial pressures, the axial growth rate was higher than that of the ethylene-grown structures, exceeding 18 mm/s at the highest powers. Surprisingly, preliminary results for the axial growth rate show that the highest normal growth rates values were obtained at the lowest pressures, as indicated in Figure 4 (top.)As for steady state diameter, exper imental results recorded in Figure 4 (bottom) indicate that increasing pres sure broadens the rod, just as for ethyl ene at partial pressures below 200 mbar. Interestingly, for all Ni(CO)4 pressures, the diameters approach a non-zero asymptotic diameter, approx imately 60 mm wide.

Interestingly, for all Ni(CO)4 pressures, the diameters approach a non-zero asymptotic diameter, approx imately 60 mm wide. One possible implication is that, for these pressures, it may not be possible to grow narrow Ni structures thinner than 100 mm, thereby limiting the resolution of the process. Further experiments at very low laser powers «100 mW) and small beam waist diameter are needed to confirm this phenomenon.Figure 5 displays measurements of length over time for a typical rod, as well as the diameter of the rod during the same time period (820 s). One can see that the axial rate rises rapidly to the steady-state growth rate (indicated NI-Fe rods cross-sections showmg the sample's by the slope of the solId lme), and then microstructure. All rods were grown at 20 mbar partial pressure at the axial rate slows to a new terminal Fe(CO)s a~d at laser P?wer of 190mWatts. The respective precursor rate, Rt.

All rods were grown at 20 mbar partial pressure at the axial rate slows to a new terminal Fe(CO)s a~d at laser P?wer of 190mWatts. The respective precursor rate, Rt. At the same time, the ro d diameter at mid-length grows steadily, beginning at 250 mm, ultimately reach ing over 1000 mm in diameter. Perhaps the most significant result from this fig ure is that, while the steady-state axial rate greatly exceeds the radial' rate (hence the rod), the terminal rate, Rt, and the radial rate are nearly the same. Consequently, for metals under a de-focused beam, the entire rod grows at a uniform rate, and the temperature gradient along the rod is small.. ....FIGURE 6.and deposit concentrations are reported below.~---,Deposit CompositionGas Mixture(I) Q.. E t1l tnNi Cone. (%)Fe Cone. (%)Ni(CO)4 Cone. (%)Fe(CO)s Cone. (%)I52.0A50.048.050.058.0B19.071.042.0,[8.055.652.0C44.4....

gas mixture composition for Auger and Microprobe measurements.3.2 NI-FE GROWTHSample structures were grown from various precursor mixtures of Nickel and Iron carbonyls. The sample composition was shown to vary with the precursor partial pressures and demonstrated the capability for alloying as well as fabricating functionally graded materials.Alloy Composition: Figure 6 shows the cross-section and microstructure of several rods grown at the same incident power, but differing precursor concentrations. The Fe(CO)5 precursor partial pressure in Figure 6.B is the lowest of the three cases (42%); the Fe(CO)5 pressure increases in the case of Figure 6.B to 50%, and reaches its maximum value at 55.6% in Figure 6.C.Throughout each sample, the ratio of Ni-Fe (in at. %) remained nearly constant due to the large capacity of the reactor chamber. The samples exhibited regions that 'Nere heavily oxidizej regions, and, in the case of Figure 6.B, inclusion of graphitic flakes.

The samples exhibited regions that 'Nere heavily oxidizej regions, and, in the case of Figure 6.B, inclusion of graphitic flakes. Yet tbe relative composition of the deposit remains constant to within ±2 percent. Figure 7 summarize~) the final composition of the Ni-Fe rods versus the gas mixture in which they were grown. The partial pressures of Ni(CO)4 and Fe(CO)5 were varied as a percentage of the total chamber pre.ssure, giving rods of similar atomic Ni-Fe ratios. In all cases, the chamber was filled with Fe( CO)5 to 20 mbar, and Ni(CO)4 was added to give a specific relative gas concentration.

In all cases, the chamber was filled with Fe( CO)5 to 20 mbar, and Ni(CO)4 was added to give a specific relative gas concentration. It is appa::ent that any Ni-Fe alloy may be produced in this manner, including high-temperature super-alloys with Ni concentrations greater than 25%.3.3 VOLUMETRIC GROWTH RATE MEASUREMENTGrowth Rate Measure: Using a stationary beam with focus at the substrate, straight-sided nee dles of graphite, nickel, iron, and nickel-iron alloys up to 7 mm in length could be deposited with out moving the laser focus. Note that this is many times the diffraction-limited Rayleigh Range of the beam focus (260 Jlm).231o.'rrt::::I=t~,e1f.n:FT~l 0.1 ~,~f:¥+'--"':,,,..;..,,..;..,,;,,....;,,,.1FIGURES.o.

The axial deposition rate, Ro, is the normal growth rate at the center of the rod, along the beam axis. The final radius of a rod, Rs, can be estimated at any axial point by integrating the component of the normal deposition rate perpendicular to the beam axis over time. The volumetric deposition rate, Rv, of a cylindrical rod depends on both the axial rate and diameter deposited over time.The normal rate, Rn, varies exponentially with laser power in the kinetic regime, while at high incident powers, it may become transport limited.Growth Regimes. Figure 8 shows the emission signature vs. time for a typical graphite rod. Figure 9 correlates the emission signature to the volumetric growth rate. After a brief delay, representative of thin film nucleation and coalescence at the substratl~, rapid axial growth commences, eventually rising to a peak rate. This interval is denoted the transient regime,2326002labeled (1) in Figure 9.

This interval is denoted the transient regime,2326002labeled (1) in Figure 9. Once the peak rate is attained, the rod grows at a constant axial rate, and it is over this steady-state regime -labeled (2) in Figure 9- that the bulk of the rod i'> grown. £ventuaIly, (at approximately 300 seconds, labeled (3) in F~gure 9) the axial ral.e slows, dropping linearly as the beam broadens and the surface temperature decreases; this is known as the tail regime.Volumetric Rate Measure. Emission spectra produced during the pyrolyt.ic decomposition of ethylene were monitored using the photodetector and filters described pre·/iollsly. The hydrogen Alpha line at 656 nm was found to be several orders of magnitude brighter than the background thermal emissions. This spectral line represents the n=3 --> n=2 transitiol~ in the Balmer Series of atomic hydrogen, as occurs following disassociation of hydrogen from C~1-:14 as well as from the thermal ionization of H2.

Since ethylene dissociates at a lower temperat:Jre than the ionization potential of hydrogen, the measured emission spectra were primarily dlle to the liberation of atomic hydrogen during pyrolysis of C2H4.Since the signal strength is proportional to the total number of disassociated hydrogen ions returning to the ground state, it was hypothesized that the emission signature is representative of the volumetric deposition rate. This was confirmed by Figure 9, where the steady-state signal amplitude is plotted vs. measured volumetric rate along lines of constant pressure. Each curve is normalized to the amplitude of the emission at incident powers of 2000 m l;V. Observe that in each case, the signal strength is linearly proportional to the volumetric rate.

Each curve is normalized to the amplitude of the emission at incident powers of 2000 m l;V. Observe that in each case, the signal strength is linearly proportional to the volumetric rate. When sufficiently well cal ibrated for the precursor partial pressure and viewing position, the signal arlplitude can be used as an absolute measure of volumetric rate.3.4 AXI-SYMMETRIC CONTROL OF 3D-LCVDThe one-to-one correspondence of volumetric rate and radius (at any given pressure) makes it pos sible to create a very stable controller for 3D-LCVD. If the spectral emission signature is fed back to the laser power controller, and compensation is made in the laser power to obtain a desired vol umetric growth rate, one obtains direct control over the axi-symmetric 3D -LCVD process.

Pro portional-derivative-integral (PID) control was used to set the laser power in real-time with feedback from the emission signature.To demonstrate, a series of samples were grown in this manner, such as the structures shown in Figure 10. All samples were grown at 200 mbar, so that the axial rate would be transport limited and to allow sufficient time for the controller to respond to the growth. (At 930 mbar, rods grow so rapidly that the steady-state regime passes in just a few seconds.) The laser power was set at max imum power to initiate growth on the graphite substrate through the transient regime. The feed back controller was then engaged after the rods had reached their steady-state radius (labelled 1 in Figure 10.A.

The feed back controller was then engaged after the rods had reached their steady-state radius (labelled 1 in Figure 10.A. The sample in Figure 10.A was grown using a linearly decreasing setpoint over the length labelled (2), while the sample in (B) was grown with an exponentially decaying total laser power in place of the PID control.The most promising feature of this control technique is that the output shape is similar to the input waveform; the linear setpoint decay in (A) resulted in a straight cone" 'Nhile the sinusoidal growth input in (C) resulted in periodic bulbs. The opposite is true for open-loop laser power con trol, where an exponential decay in the power is required to produce a linearly tapered rod.2333.5 GRAPHITE TEMPERATURE MEASUREMENTSTemperature Gradient Measurements: In this section, the temperaturf.; gradient over sever,tl graphite rods will be measured from the blackbody radiation emitted during their growth. Kodak High Speed infrared emulsion no.

Kodak High Speed infrared emulsion no. 2681 was employed throughout the experiments. A standard 250 mm telephoto lens was used in combination with a 250 mm achromat, achieving a 1.2x mag nification of the object on the film. Exposures were made at various wavelengths using narrow- band filters.Temperature Measurements: Figure 11 displays several successive photographs of the same rod taken at various wavelengths; the frames were taken during the steady-stare growth regime in rapid succession. The rods were viewed at a 45±1· angle from the laser ax!:;. The resolution of the photographs is approximately 7 Jlm, and the direction of the incident beam is shown by the dashed, blue arrows. The brightest regions are the surface of the rod; these are color-coded red, orange, and yellow, in descending intensity according to the color bar.

The brightest regions are the surface of the rod; these are color-coded red, orange, and yellow, in descending intensity according to the color bar. The gas surrounding the rods also emits weakly, and this appears primarily as green, blue, and purple surrounding the brighter regions.234Temperature measurement at various wavelengths during rod growth.Finally for the rod grown at 2930 mW, the power is sufficient at the tip to reach graphite's sub limation temperature of 3925 K; in fact, SEM photos of this rod exhibit s:ight dimpling, such as found more prominently on rods grown at higher laser powers-and previously attributed to mass transport limitation. A slight warpage near the rod tip is also visible, possibly due to softening of the graphite at excessive temperature.

A slight warpage near the rod tip is also visible, possibly due to softening of the graphite at excessive temperature. It remains to be seen if the dimple effect is entirely caused by diffusion limitation, sublimation, or a combination of both phenomena.4 CONCLUSIONS Controlled 3-dimensionallaser-induced vapor deposition of Ni-Fe alloys is feasible from nickel and iron carbonyls, producing alloys with less than ±2 at.% compositional variance. Sustainable steady-state growth rates of up to 30Jlm/s are possible.

Sustainable steady-state growth rates of up to 30Jlm/s are possible. Further studies will determine the detailed microstructure and hardness of the Ni-Fe deposits, as well as the activation energy of Fe(CO)s.235In addition to controlling the laser power, the emission signature could also be used for focal positioning, comparing the actual emission amplitudes with the expected steady-state rates, com pensating to maintain the steady-state focal position.Direct feedback control of the growth rate will also be essential for layered growth, as previ ous attempts to build-up shapes by repetitive scanning have proven to be unstable. As one layer is grown over another using open-loop control, perturbations in the thickness of the underlying deposit induce exponential perturbations in the next layer (kinetic regime), yielding irregular deposits.

With growth rate control, compensation can be made for such perturbations, and rate controlled 3D-LCVD can be used for both layered and continuous growth.Similarly, the growth control method could be used to obtain consistent direct writing resultsover varying topographies and substrate materials.5 ACKNOWLEDGEMENTSThe authors wish to thank the National Science Foundation (Grants: NSF ECS-9314071 and NSF DDM-9057059) for their generous support. Grateful acknowledgments also go to Ray Dove of the Rensselaer MRC microscopy laboratory for the SEM photos, Dr. Hudson of the Rensselaer Materials Engineering Department, for use of the Auger Microprobe, Nonn Gendron, of Rensse laer, for advice in sectioning the samples, and David Wark of the Rensselaer Geology Dept. for contributing to the x-ray microprobe analysis. Thanks finally to the CIEEM staff for helping to dispose of the iron and nickel carbonyls.6 REFERENCES [Bauerle, 1984] D.

RECT FREEFORM FABRICATION OF SPATIALLY HETEROGENEOUS LIVING CELL-IMPREGNATED IMPLANTSD. L. Cohen, E. Malone, H. Lipson, and L. J. Bonassar Department of Mechanical and Aerospace Engineering, Cornell University, Ithaca, NY 14850 Reviewed, accepted August 4, 2004ABSTRACTThe objectives of this work are the development of the processes, materials, and tooling to directly “3-D print” living, pre-seeded, patient-specific implants of spatially heterogeneous compositions. The research presented herein attempts to overcome some of the challenges to scaffolding, such as the difficulty of producing spatially heterogeneous implants that require varied seeding densities and/or cell-type distributions. In the proposed approach, living implants are fabricated by the layer-wise deposition of pre-cell-seeded alginate hydrogel.

In the proposed approach, living implants are fabricated by the layer-wise deposition of pre-cell-seeded alginate hydrogel. Although alginate hydrogels have been previously used to mold living implants, the properties of the alginate formulations used for molding were not suitable for 3-D printing. In addition to changing the formulation to make the alginate hydrogels “printable,” we developed a robotic hydrogel deposition system and supporting CAD software to deposit the gel in arbitrary geometries. We demonstrated this technology’s capabilities by printing alginate gel implants of multiple materials with various spatial heterogeneities, including, implants with completely embedded material clusters. The process was determined to be both viable (94±5% n=15) and sterile (less than one bacterium per 0.9 µL after 8 days of incubation). Additionally, we demonstrated the printing of a meniscus cartilage-shaped gel generated directly from a CT Scan.

Additionally, we demonstrated the printing of a meniscus cartilage-shaped gel generated directly from a CT Scan. The proposed approach may hold advantages over other tissue printing efforts [5,9]. This technology has the potential to overcome challenges to scaffolding and could enable the efficient fabrication of spatially heterogeneous, patient-specific, living implants.INTRODUCTIONThe aim of this research is to develop a cell-seeded gel that is compatible both with solid-freeform fabrication and with biological cell. Alginate gels have been previously used for non SFF processes such as injection molding of living implants. Alginate is a seaweed derivate which, when combined with certain calcium cross-linking compounds, forms a polymer hydrogel [1]. Cells can be evenly mixed into this gel and then injection molded in order to create a desired geometry [3].

Cells can be evenly mixed into this gel and then injection molded in order to create a desired geometry [3]. When the cell-seeded gel is placed in a growth medium and incubated, the cells consume the medium and replace the alginate gel, resulting in living tissue. The material properties of the gel are governed by the type of alginate, the type of cross-linker, the concentrations of each, as well as other factors such as temperature and age. The material properties that are optimal for molding the gels are not the same as those required for printing the gels. The desired material properties for both cases are essentially opposite of each other; the discrepancies between the material properties required for molding and printing pose a major design challenge that this project addresses. The type of cell that is seeded into the gel can also dictate the required material properties.

The type of cell that is seeded into the gel can also dictate the required material properties. In this research, bovine cartilage chondrocyte cells were used because they do not require a blood supply for nutrition, but the challenge of growing cells that do require a blood supply is a separate open research field. Another consideration when dealing with alginate gels is the need for sterility. Much in the same way that the alginate gel supports growth of seeded cells, it will also support the growth of bacteria and fungi; hence, sterile print conditions are essential to prevent infection.720GEL PRINTING SYSTEMRobotic Test PlatformWe have designed an open-architecture gantry robot that was used in this research [4]. The tool paths are generated by our path planning software takes multiple stereo lithography (STL) files as input, corresponding to multiple material types.

The tool paths are generated by our path planning software takes multiple stereo lithography (STL) files as input, corresponding to multiple material types. This system has been designed to allow the printing of multiple materials within a single part; in this research, this capability is used to print multiple gels with varying cell types, chemical concentrations, cell densities, etc. The robot is capable of moving the deposition tool to a specified position with accuracy and repeatability of ±25 µm. The full specifications of this robotic system are given in Table 1.Table 1 - Fabrication System Performance SpecificationsMinimum material stream/drop diameter Materials cross-section area Build rate Nominal speed along path Min. turn radius at nominal speed Tool position accuracy (±) Tool position repeatability (±) Positioning resolution Build envelope x Build envelope y Build envelope z Max.

Additionally, the tool has to enable efficient material changing. Efficient swapping of sterile materials is important for printing implants of multiple gels (each with different cell types, concentrations, etc.).Through experimentation, we found that elastic materials are difficult to deposit with pneumatic dispensing systems and instead a volumetrically controlled dispensing system was chosen [4]. An ABS plastic rapid prototyped frame connects a linear actuator to the syringe cartridge (Fig. 3). Medical-quality sterile syringes are used as the disposable material cartridges of the deposition tool. Luer lock syringes were selected in order to enable a wide variety of syringe tips to be utilized.

Luer lock syringes were selected in order to enable a wide variety of syringe tips to be utilized. Tool performance specification can be found in table 2.Table 2 – Deposition Tool Performance SpecificationsMaximum applied pressure1592 kPaCartridge volume Maximum volumetric flow rate10 mL 10.6 mL/sDeposition accuracy0.000426 mLDeposition precision0.000426 mLFigure 3 - CAD Model of Rapid Prototyped (Linear Deposition Tool Components Actuato and Syringe Not Shown)Figure 4 - Picture of Deposition Tool with (Mounted To Disposable Syringe Attached Robot’s Tool Platform)722Sterile Printing EnvelopeSterile printing conditions are a paramount concern, the lack of which would allow the printed implants to become infected.

Although, as described above, a solution was found for keeping the material sterile during deposition, a way of keeping the printed pieces’ immediate environment sterile still needed to be found.One possible solution to this challenge was to build a sterile, hermetically sealed envelope around the immediate printing zone. While a full envelope that encompassed the printer was considered, an even smaller envelope that surrounds only the printed piece was more desirable.A 1-millimeter thick, autoclavable, clear plastic bag was chosen as the envelope. This bag could be loaded with a Pyrex Petri dish, sealed, and then autoclaved to ensure that the entire inside of the bag, including the Petri dish, was sterile. At this point, the inside of the bag served as a closed, sterile environment (Fig. 5). At the time of printing, the tip of the syringe and a section of the outside surface of the bag were swabbed with 70% ethanol solution and then the syringe tip was inserted through the bag.

5). At the time of printing, the tip of the syringe and a section of the outside surface of the bag were swabbed with 70% ethanol solution and then the syringe tip was inserted through the bag. All deposited material was printed onto the sterilized Petri dish and surrounded by the previously autoclaved environment. After printing, the bag was taken to a sterile work hood where the outside of the bag was sprayed with alcohol and the sample was removed (Fig. 6). The disposable autoclave bag could then be discarded and a new bag used for the next print.Figure 5 – Top View of Sterile Printing EnvelopeFigure 6 - Sterile Sample Being Removed From Envelope Under Ultra Violet (UV) Sterilized HoodPRINTABLE GEL FORMULATIONThe gel is composed of two components: a solution of alginate and a calcium cross- linker solution.

Even though there are only two components, there were many experimental variables and it was a challenge to find a gel that was both supportive of cellular life and “printable.”723Formulation ConstraintsThe alginate gel had to be both suitable from biological and SFF standpoints. In order fora gel to be considered biologically suitable, the alginate gel had to sustain cellular life.In order to satisfy the SFF needs, the gel had to be “printable.” By “printable,” it is meant that the gel would: 1) bond between layers, 2) hold its shape against gravity, and 3) cross-link slowly enough to prevent cell and material shearing during printing.Formulation ProcessThe first step was the identification of experimental parameters.

These parameters were identified as: 1) the type of cross-linker, 2) the cross-linker concentration, 3) the molecular weight of the alginate 4) the M/G ratio1 of the alginate, 5) the concentration of the alginate, 6), how long to mix the alginate and cross-linker together and with what force 7), how long to let the gel cross-link before printing, 8) the choice of additives, and 9) the range of syringe tip diameters that can be used to print the gel.Since two separate sets of needs, biological and SFF, had to be satisfied by some single combination of the above nine parameters, each parameter’s effect on the biological and SFF qualities of the materials had to be analyzed.Biological Formulation NeedsFrom prior publications [10], it was clear that the gel would most likely be viable, i.e., support cellular life, as long as the alginate concentration was 2% in PBS2 and the cross-linker concentration was less than or equal to 2% in PBS.

A standard seeding density of 50 million cells per milliliter of gel was also used. These established standards gave reason to believe that as long as the standards were adhered to, the biological needs would be satisfied. Also, the standards served to constrain two parameters, hence they simplified the formulation process by eliminating two variables.SFF Formulation NeedsWhile the biological needs were satisfied by the aforementioned published standards, the SFF-related needs had yet to be satisfied.

Extensive experimentation ensued in order to determine what effects the nine parameters had on the three SFF-related material properties: 1) bonding between layers, 2) resistance to weight induced deformation, and 3) prevention of shearing due to rapid cross-linking.1 M/G ratio is the mannuronic acid to guluronic acid ratio 2 PBS (Phosphate-Buffered Saline) is an isotonic solution that is commonly used as a dilution agent when making hydrogels724Inter-layer Bonding ExperimentsTo study the effects of the nine parameters on the first SFF-related need, bonding between layers, it was determined that this behavior was directly related to the rate and amount of cross-linking at the time of printing. Gels that cross-linked more quickly would be further cross-linked at the time of deposition and would be less receptive to subsequent layers because the bonds were already formed before the following layers could be deposited.

Experiments were run in which the experimental parameters were varied and the cross-linking rate was given a relative, qualitative measure depending on how quickly the gel stiffened. The relationships between the experimental parameters and the ability of the gel to bond between layers are summarized in Table 3.

The relationships between the experimental parameters and the ability of the gel to bond between layers are summarized in Table 3. The goal of this experiment was to find some combination of the nine parameters that yielded a gel which cross-linked as quickly as possible, yet still bonded between layers.Table 3 - Relationship Between Parameters and Successful Interlayer BondingExperimental ParameterRelationship with Successful Bonding Between LayersIncreased alginate molecular weight Increased cross-linker concentration Longer mixing time Longer sitting time before printing Larger tip diameter(-)(-)(-) (-)(+)Resistance to Weight-Induced Deformation ExperimentsIn order to see what effect each parameter had on the second SFF-related need, resistance to weight deformation, each parameter’s effect on the gel viscosity was studied using an automatic viscometer.

It was initially thought that the viscosity would be an accurate indication of how well the gel would hold its shape against gravity. In order to further vary the viscosity, dextrose was used as a bulking agent. However, after comparing the results of the viscosity experiments to simulated prints in which the gel was layered by hand, it was clear that viscosity was not an accurate indicator of a gel’s resistance to weight-induced deformation. The notion was rejected and this part of the formulation process was iterated. At this point, it was noticed that the resistance was strongly related to cross-linking rate. While a high viscosity did not help the gel hold its shape, partial cross-linking did; therefore, the faster a gel cross-linked the better it would be able to hold its own weight. At this point, it was realized that two SFF-related needs were directly determined by cross-linking rate, that is, the bonding between layers and the resistance to gravity.

At this point, it was realized that two SFF-related needs were directly determined by cross-linking rate, that is, the bonding between layers and the resistance to gravity. However, these two needs were opposed to each other. A higher cross- linking rate would help the gel hold its shape against gravity, while at the same time it would make the layers less likely to bond with each other.

Some acceptable compromise needed to be reached.725Table 4 - Relationship Between Parameters and Resistance to Weight-Induced DeformationExperimental ParameterRelationship with Resistance to Weight-Induced DeformationIncreased alginate molecular weight(+)Increased cross-linker concentration(+)Longer mixing time Longer sitting time before printing Larger tip diameter(+) (+) (+)Table 5 - Effect of Cross-linking Rate on SFF NeedsSFF-Related Need Bonding Between LayersResistance to Gravity Prevention of ShearingHigh Cross-linking Rate Layers will not bond with each other (-) Gel can hold shape (+) Gel will shear and lose mechanical properties (-)Low Cross-linking Rate Layers will bond with each other (+) Gel cannot hold shape (-) Gel will not shear (+)Cross-linking Rate ExperimentsThe third SFF-related need was also directly related to cross-linking rate; this was prevention of material shearing due to rapid cross-linking.

If the material cross-linked too quickly, then the gel would significantly bond before being deposited and the bonds would be sheared upon passing through the constricted syringe tip. A summary of the relationship between cross-linking rates and the SFF-related needs can be found in Table 5.Manual Test PrintsIn order to find a cross-linker type/concentration that was suitable for the three SFF formulation constraints, various chemical combinations were tested manually (fig. 7). An efficient and simple test that indicates a particular formulation’s printability is a conical stacking experiment. One half inch to one inch cones are deposited manually over a time range from t = 0 to t = 20 minutes. Each stacks ability to hold its shape against gravity is observed. Since the material properties are time dependent, testing over a large time period is important. Not only are qualitative traits noted for each formulation, but also an optimal time frame for each formulation is identified.

Not only are qualitative traits noted for each formulation, but also an optimal time frame for each formulation is identified. In this optimal time frame, the properties are suitable for printing. With some formulations the optimal frame begins at t = 0, but with others the optimal time window begins after a delay. The goal was to find a printable formulation with longest optimal time window.Selected FormulationAlthough we regularly use gel formulations of various cross-linkers, alginate-types, and mixing styles, the formulation used in the experiments reported herein is 2% Protanal Alginate (a high M-group alginate) in PBS and 0.5% Calcium Sulfate in PBS. The two solutions were mixed with each other in a 2:1 alginate to CaSO4 ratio. Since the mixing technique greatly726affects the resulting gel, due to shear effects, it is important to mix the gel consistently.

Since the mixing technique greatly726affects the resulting gel, due to shear effects, it is important to mix the gel consistently. The mixing technique that we used was to take one 10 mL Luer lock syringe with alginate, another with cross-linker and mix the two back-and-forth through a Luer lock stopcock. We found a successful mixing procedure to be 10 full-cycle pumps at a rate of 1 pump per second. In consideration of resolution and shearing constraints, a 0.030” diameter syringe tip was chosen.Figure 7 – Manual Print Simulation for Testing Material PropertiesGEL PRINTING EXPERIMENTSUnseeded Gel Printing TestThis experiment verified the ability of this system to print gels of complex arbitrary geometries. In this experiment, various geometries were printed on the gantry robot in full-scale, non-sterile printing runs.

In this experiment, various geometries were printed on the gantry robot in full-scale, non-sterile printing runs. The following “proof of concept test pieces” were chosen because they each demonstrate the unique capabilities of this technology and because they have interesting medical applications.Viability of Seeded/Printed Gel TestThe viability test verified short-term cellular survivability. In other words, the test determined whether the gel sustained life throughout the printing process, i.e., while being subjected to the shear forces induced by the deposition process. During viability testing, the printing process was simulated and then the percentage of living cells versus dead cells was measured before and after the simulated printing process (Fig. 11). In order to count the percentage of living cells, two fluorescent chemical markers were used. Calcein acetoxymethyl (calcein AM) attaches to living cells and ethidium homodimer-1 (EthD-1) attaches to dead cells.

Calcein acetoxymethyl (calcein AM) attaches to living cells and ethidium homodimer-1 (EthD-1) attaches to dead cells. Through the use of a fluorescent light source, a microscope, and the appropriate optical filters, the living and dead cells were identified (figs. 13, 14).

The viability test did not have to be done under sterile conditions; the test was measuring short-term viability and infection would only have a long- term effect.727Figure 8 - Test Piece 1: Meniscus-Shaped Single Material Gel (Above Are the CADFigure 9 - Test Piece 2: Two Gel Embedded Piece (Red Dyed Gel Embedded in ClearFigure 10 - Test Piece 3: Intervertibral Disk-Shaped, Two Material, Spatially728Figure 11 - Isometric View of Deposition Too and Printer Before Printing Into a Petri Dishl Figure 12 - Side View of Deposition Tool DuringExperimentFigure 13 – Microscope View During Viability Test using Fluorescence Microscopy (Filtered to illuminate living cells)Figure 14 – Microscope View During Viability est using Fluorescence Microscopy (Filtered to illuminate dead cells)TProcess Sterility TestThe approach of using an autoclavable plastic bag as a printing envelope had to be tested to erify that no bacteria or fungi were allowed to infect the printed alginate gel.

A sterile sample v of alginate gel was prepared and placed in a sterile syringe. This syringe was then transported, with a sterilized syringe end-cap, to the printer and the full printing procedure was executed. After the printing, the sample was brought back to a sterile work hood where the envelope was opened and the sample removed. The gel sample was incubated for a period of eight days. After this incubation period, the sample was analyzed for presence of bacteria and fungi growth with the use of a chemical marker and fluorescence microscopy. It was determined that there was less than 1 bacterium per 0.9 µL after 8 days of incubation; hence, the bag envelope concept was proven to be a success.729Direct CT Scan Printing TestThis research additionally demonstrated the ability to efficiently print a hydrogel implant directly from a CT Scan.

A CT Scan of a meniscus cartilage was provided by the Cornell affiliated H ospital for Special Surgery (HSS) and then converted into an STL file. The STL file was then loaded into the CCSL software package and prepared for printing. The results of this direct CT Scan print are shown below.Figure 15 - Meniscus-Shaped Piece from CT Scan During Printing (above) and CAD Model Used for Print (below)Figure 16 - Mensicus-Shaped Piece After Printing (above) and CAD Model Used for Print (below)730CONCLUSIONSe have demonstrated the ability to successfully 3-D print viable, infection-free, living hydrogel W le materials with complex biological geometries. By printing with alginate implants of multip ydrogels pre-seeded with gels, this technology serves as a versatile tissue engineering platform h on which spatially heterogeneous implants of varied seeding densities, chemical concentrations, cell-types, etc. can be fabricated in arbitrary geometries.

rect Generation of Contour Files from Constructive Solid Geometry RepresentationsSashidhar Guduri, Research Assistant Richard H. Crawford, Assistant Professor Joseph J. Beaman, Professor Department of Mechanical Engineering The University of Texas at Austin Austin, TX 78712AbstractGeometry processing for layer-based Solid Freeform Fabrication consists of at least two steps: slicing the geometry to obtain the part contours for each layer, and scan-converting the layers for laser scanning (or other device-dependent in-layer processing). This paper discusses the generation of contour files directly from Constructive Solid Geometry (CSG) representations for the Selective Laser Sintering process. Previous work at The University of Texas focused on slicing CSG representations composed of quadric primitives. This paper extends previous work at UT to include the torus, a fourth degree surface, as one of the CSG primitives.

This paper extends previous work at UT to include the torus, a fourth degree surface, as one of the CSG primitives. Slicing a torus results in a fourth degree equation in two variables, which represents a curve in two-dimensional real space. For. some special cases, this fourth degree equation may be sub-divided into two second degree equations. For the cases where the fourth degree equation cannot be sub-divided, a method is presented to approximate the fourth degree curve with second degree curve segments.IntroductionSolid Freeform Fabrication (SFF) techniques manufacture solid objects directly from three-dimensional computer models. Most SFF processes produce parts on a layer-by-Iayer basis. process begins by slicing the geometric description of the part into layers. The slicing operation generates the contours of the part for each layer. The contours are then processed in a manner dependent upon the particular SFF technology.

The slicing operation generates the contours of the part for each layer. The contours are then processed in a manner dependent upon the particular SFF technology. For instance, for Selective Laser Sintering (SLS) the contours are discretized into "toggle points" at which the laser beam must be modulated to produce the desired solid.geometric description used to represent solid objects significantly affects the accuracy and quality of the final parts produce with SFF. One way to improve the final accuracy and definition of SFF parts is to improve the. geometric descriptions that represent three dimensional objects. As described in [1], ConstructiveSQlidGeometry is one geometric description where the accuracy iofthe final contours can be improved as compared to present geometric representations. A method was presented in [1] for generating contour files from Constructive Solid Geometry representations composed of quadric objects.

A method was presented in [1] for generating contour files from Constructive Solid Geometry representations composed of quadric objects. This paper extends that work to include higher degree surfaces (degree greater than two) in the primitive set. Special attention is given to the torus, a fourth degree surface.Slicing a higher degree surface results in a boundary curve whose degree may be greater than two. A curvec)f degree .greater than two is parametrizable if the genus (g) of the curve is zero. The genus of a curve is defined by291g = (N -l)~N - 2) _ ~O(p),where N is the degree oftheeq1Jation and the O(Pi) operator appropriately countsthe number of times that the curve comes in contact with itstelf at each singular point Pi' All quadratic curves are genus zero and are therefore parametri~able. A limited set of curves of degree greater than two are parametrizable.

A limited set of curves of degree greater than two are parametrizable. Unfortunately, the algebraic equations produced by many geometric design applications are not generally genus zero. It therefore becomes necessary to approximate higher degree curves with lower degree curve segments. These curve segments are then parametrized individually.The method presented in this paper approximates curves of degree greater than two using second degree curve segments. There are two advantages to using second degree curves. First, all second degree curve segments are parametrizable. Second, the intersection of two second degree curves (required for Boolean set operations) can be computed by well-known, stable algorithms. If curve segments of degree greater than two are used as approximations, calculating the intersection of two curves becomes computationally expensive. The algorithm for generating the approximation is based on degree reduction of the triangular Bernstein form of the curve.

The algorithm for generating the approximation is based on degree reduction of the triangular Bernstein form of the curve. The next section of the paper illustrates the representation of the torus as a CSG primitive and gives the equations involved in calculation of the implicit equation of the torus. Following this section, a method for determining the Bernstein coefficients of the implicit curve is described. The remainder of the paper describes the steps in the approximation algorithm: parametrization, approximation and error estimation, subdivision, and resolution of singularities.

The remainder of the paper describes the steps in the approximation algorithm: parametrization, approximation and error estimation, subdivision, and resolution of singularities. Examples are presented at the end of the paper.Torus as a CSG PrimitiveThe equation of a torus generated by sweeping a circle about the y axis, with the origin atthe center, is given by:As a CSG primitive the torus is represented using a base point, an axis vector and two radii, as illustrated in Figure 1.Given the above information, the equation of the torus can be generated by translating the base point of the torus to the origin and rotating about x and z axes such that the resulting axis of the torus is aligned with the y axis. This procedure is summarized below:1. Normalize the axis vector of the torus. 2. Translate the base of the torus to the origin.3. Rotate about the x axis by an angle of -tan-1 (;:) • 4. Rotate about the z axis by an angle of tan-1(..J x. J.

2. Translate the base of the torus to the origin.3. Rotate about the x axis by an angle of -tan-1 (;:) • 4. Rotate about the z axis by an angle of tan-1(..J x. J. z; +y;If the above steps are applied the following transformations are obtained292(1)(2)xFigure 1. Representation of a torus.z =--=====1)(3)where into equation 2 gives the fourth degree equation of the torus value of the slicing plane gives a fourth degree equation in the contour for that slicing plane.and Z define the coordinate system local totorus. Substitutingequations z equation ofx, y and z. Substituting y that is the'n.llI'1I'llTQlI'01I::'1I11'1II"lI of Algebraic EQualUOllS tosection describes a method for """....,,,""'...t-...., ,...triangular regions .n""",...."""" elevation of u..l.f'O.""..,.l.u..I."",.".,,...........

Finally the Bernstein coefficients are given by:N b.. =I)D~I)P(N,i,j)After the Bernstein coefficients are obtained, the algorithm proceeds with approximation, as described in the next section.Approximation of Higher Degree CurvesThe algorithm for approximating algebraic curves of degree greater than two consists of four stages [3]. The parametrization stage builds a parametric definition. The approximation stage attempts to build a degree N-l equation that matches the given degree N expression within acceptable error. The subdivision stage partitions the region where the curve is defined into smaller regions, if the region is not already too small. In the fourth stage, the algorithm resolves singularities.

The subdivision stage partitions the region where the curve is defined into smaller regions, if the region is not already too small. In the fourth stage, the algorithm resolves singularities. The complete algorithm is summarized as follows:Stage 1: ParametrizationAn attempt is made to parametrize the curve f(x, y) = °if the degree is less than 3.Should the degree be greater than 2, the algorithm proceeds with the approximation stage, where the degree of the equation is reduced to two.Stage 2: ApproximationIf the degree k function f(x, y) monotonically increases or decreases with respect to any one edge of the domain triangle, estimate the error present in the k-l approximation. If the error is within limits, return the approximation to stage 1 of the algorithm; if not, proceed with the subdivision stage.Stage 3: SubdivisionIf the size of the triangular region is smaller than a preset limit, assume the region contains a singularity and proceed to stage 4.

The intersection points of the curve and the Bernstein triangle are calculated and the appropriate curve segments are taken by determining if the curve segment lies inside the triangle or not.ApproximationIn this stage of the algorithm approximations of degree N-1 are generated for degree N curves and the characteristics such as convex hull property of the Bernstein polynomial basis are exploited to estimate the maximum error present in the approximation. This section describes the equations involved in degree reducing and degree elevating Bernstein polynomials.Given a degree N algebraic curve in Bernstein form, an exact representation of this curve can be created using a degree N+1 Bernstein polynomial basis [5]. Mathematically, this means the expression+'i+j::;(N+l)L .,.'(N 1-: - ') 1.J. +1 Ji,j=Ois equivalent toN'i+j::;NL .,. '(N -'. - ') J i,j=O 1.J.1This degree augmentation process, shown in the following equations,.

k =--~--~--:.-_---=-i1- ,j,i :;t: 1 and i ~ j, kN*l~ . 1 k =--~---~----"'----j1,]- ,j :;t: 1 and j ~ i, kN*e. k 1. =--~-----.;~--~~kl,j,k:;t: 1 and k ~ i,jUnlike degree elevation, the coefficients in degree reduction are defined only in terms of previously computed coefficients, and all coefficients with negative subscripts are zero. When any two or all three indices are equal, the components given by each of the corresponding equations are averaged to define a single value. Note that if the degree of a polynomial is first subsequently lowered, the original polynomial coefficients are retrieved. elevatedOnce the degree reduction procedure has been applied to the algebraic curve f(x, y) =0 and an approximation g(x, y) is generated, an estimate of the approximation error is required.

elevatedOnce the degree reduction procedure has been applied to the algebraic curve f(x, y) =0 and an approximation g(x, y) is generated, an estimate of the approximation error is required. An upper bound on the error can be derived from a combination of difference values (dz = f g) and directional derivative information taken from the two single valued surface equations z = f(x,y) and z =g(x,y) [7].To begin, the degree elevation procedure is invoked to ensure f and g have the same polynomial degree, thereby assuring the same control point lattice over the triangle. Applying the convex hull property, the maximum difference between the two surfaces z =f(x,y) and z = g(x,y) is bounded by the largest difference found between the two control nets:Tangent plane and normal information can be derived from the surface equations [5, 8]. This information is used to relate the largest surface difference, given in the above equation, to the difference between the two algebraic curves.

This information is used to relate the largest surface difference, given in the above equation, to the difference between the two algebraic curves. If point P is located on the curve approximation g =0, the corresponding point on f =0 has to be located. Figure 3 illustrates that there must exist297(15)(16)(17)........- - - - - - - -1.......... "sLlC9 =0Figure 3. Error estimation.a point on f = 0 at least within a distance i1C, where S is the smallest angle between the tangent plane and a direction s defined in the x-y plane, and i1C is given by the equationi1C = i1z tanSSince the surfaces are single valued, tanS defines the value of the directional derivative, ~~, of the function z = f(x,y) with respect to the direction s.

If the original Bernstein triangle with its Bernstein coefficients is called the original patch, then the Bernstein coefficients of the directional derivative patch can be calculated from the original patch using the following equation (for direction s shown in Figure 4):b.. = N * (B. . 1 - B.. ) ,I,)1,)+1,)where i+j <= (N-I) and N is the degree of the original patch.The convex hull property is applied to the directional derivative patches of f(x,y) and g(x,y) to determine a minimum value for tanS. Combining the maximum surface difference withthe minimum value of tanS, a single error bound is produced:298(18)(19)(20)Figure 4. Original patch and its directional derivative patch.tighter error bound, but computationally more expensive, can be obtained by calculating the the control points. If the error estimate is not within the specified limits, the triangleerror at must be subdivided.

If the error estimate is not within the specified limits, the triangleerror at must be subdivided. Subdivision techniques are discussed in the next section.SubdivisionThe subdivision stage of the algorithm is reached under two sets of circumstances. When the given function is not monotonic with respect to any edge of the triangle, the triangle is subdivided to generate new sets of edges. The second condition is reached when an unacceptable approximation error is present. The following equations define the procedure for calculating Bernstein coefficients after subdivision [7]:b?ok(u,v,w)=boOk'l,j,l,j,where i+j+k =N, andwhere i+j+k =N-m.In general, sets of control points from a degree k control net are combined three at a time, weighted by the barycentric coordinates, to generate a degree k-l control neLThisprocess is repeatedN times for a degree N Bernstein surface, ultimately producing a single. value which is the desired point on the surface.

value which is the desired point on the surface. The three sets of vertices b~,j,k(u, v,w), bl,o,k(u,v,w) and btj,o(u, v, w) define the control nets for the new surface patches generated by subdividing the surface at the point with barycentric coordinates (u,v,w).The geometric interpretation of the recursion process is shown in Figure 5.Resolution of SingularitiesA singular point is a point on the curve where the function and its first partial derivatives vanish. One may assume that a region contains a singularity if the side of the triangle becomes smaller than the approximation tolerance. When a region is discovered to be near a singularity, the following quadratic transformation is used:299(21)(22)bO1,0,2bO2,0,1bO3,0,0Figure 5. Subdivision of a bivariate Bernstein polynomial.x=xA piecewise linear approximation is generated near singularities by numerically marching along the various branches of the proper transform.

Subdivision of a bivariate Bernstein polynomial.x=xA piecewise linear approximation is generated near singularities by numerically marching along the various branches of the proper transform. The sequence of points generated are then mapped back to the original coordinate system by reversing the quadratic transformation.ExamplesExamples are presented in this section demonstrating the approximation of a cross section of a torus using a collection of second degree curves. The curve traces were generated using a collection of procedures written in the C programming language and executed on a Sun™ SparcStation 2. In each of the examples, approximation is developed over a triangular region enclosing the closed curve.The torus used for slicing is centered at (0, 0, 0) and the axis of the torus is straight line given by the equation y =x.

The radius of the torus is 3, and the radius of the disc rotated about the axis to generate the torus is 0.7.Figure 6 shows the cross-section of the torus at z =2.2. The number of quadratic curvesegments in this approximation is 306.300(23)~ContoursFigure 6. Torus sliced at z = 2.2.Figure 7 shows the cross-section of the torus at z =2.3 (the singularity case). The numberof quadratic curve segments in this approximation is 318.~Contours.Figure 7. Torus sliced at z =2.3.ConclusionsThis paper discusses a method to directly process Constructive Solid Geometry representations and obtain contour files. Aspects involved in slicing higher degree surfaces, in particular the torus, are highlighted. The algorithm is applicable to other higher degree surfaces301as well, including rational bicubic parametric surfaces. The algorithm provides a rational basis for approximating geometry for SFF applications.References1. Sashidhar, Guduri., Crawford, RichardH.

rect Injection Molding Tooling Inserts from theProcess with Copper PolyamideChristian Nelson, Jason Kepler, Rick Booth, Phillip ConnerDTM Corporation, Austin,The "RapidTool" Short Run (SR) Tooling Process using the Copper Polyamide material provides a route to mold inserts for injection molding made directly in the Selective Laser Sintering machine. The STL files for the mold inserts are shelled and conformal cooling lines and ejector pin guides are added before SLS processing. in the SLS machine provides quick metal/plastic tooling with good thermal conductivity. Final preparation of the tooling inserts includes sealing the surface with epoxy, final finishing using sandpaper, and backing up the shells with a metal alloy.

Final preparation of the tooling inserts includes sealing the surface with epoxy, final finishing using sandpaper, and backing up the shells with a metal alloy. The Copper Polyamide SR Tooling inserts are used to mold several hundred parts with common plastics with injection cycle times similar to conventional molding cycle times.Sintering of the materialRapid Prototyping and Rapid Tooling continue to change how products are designed and tested. Prototyping equipment used to make plastic models for visualization ten years ago is now used to· make investment casting patterns, sand casting cores and molds, patterns for secondary processes, functional prototypes, and mold inserts for injection molding.Market acceptance of Rapid Tooling continues to grow as designers realize the benefits of molding prototype parts with the actual production material. DTM introduced RapidSteel 1.0 in 1996 for the creation of steel/copper mold inserts [1].

DTM introduced RapidSteel 1.0 in 1996 for the creation of steel/copper mold inserts [1]. DTM introduced the next generation of metal material in 1998. The RapidSteel 2.0 material provides improvements in processing time, finishing time, and accuracy compared to the original RapidSteel material. 3D Keltool was introduced by 3D Systems as a secondary process to produce steel/copper inserts from a RP pattern [2]. Extrude Hone announced the ProMetal RTS-300 system for creating porous steel inserts using the three-dimensional printing (3DP) technology licensed from MIT [3]. These processes require a furnace to sinter the steel and infiltrate the porous steel part with a copper alloy. The resulting parts are fully dense and can be used to mold over 100,000 plastic parts with most plastics.Recently, a new breed of Rapid Tooling solutions has emerged to provide short runs of production equivalent plastic parts.

These limited runs of molded plastic parts are often used in the product development stage where several hundred parts in the final material are needed. The faster and more economically a manufacturer can get a supply of parts produced in actual production materials, the faster he can perform life testing, make adjustments to the design if necessary, and get the new product on the market.Methods of creating short run mold inserts available in today's market include direct manufacture using RP, cast metal/epoxy tooling, and CNC machining. Mold inserts made directly in rapid prototyping devices include SLA epoxy, SLS metal/plastic, and SLS direct metals [4]. The metal/plastic material available for processing in the Selective Laser Sintering process is Copper Polyamide introduced in mid 1998.451The Copper Polyamide material is a heat resistant, thermally conductive composite of copper and plastic that was introduced for use with DTM's Sinterstation System.

The Copper PA material processes in the Sinterstation at conditions similar to the DuraForm Polyamide material. The Sinterstation builds the mold inserts directly from the geometry described in the STL file. Features like runners, gates, conformal cooling lines, and ejector pin guides are included in the STL file and built directly into the part. Turnaround times for the production of a mold insert are as short as a day. The key is that no furnace process is involved.The Copper PA composite makes parts that are machinable and easily finished with wet sanding. Heat resistance and thermal conductivity are better than most plastic tooling materials, and it is possible to mold parts with cycle times that approach production rates.

Heat resistance and thermal conductivity are better than most plastic tooling materials, and it is possible to mold parts with cycle times that approach production rates. Injection mold inserts made with this material are used to mold 100 to 400 parts in polystyrene (PS), polyethylene (PE), polypropylene (PP), glass filled polypropylene (GF PP), ABS, PC/ABS, and other common plastics.Mold PreparationInjection molding incorporates a cavity into which hot plasticized material is injected under pressure. Heat is removed from the material in the mold until it is rigid enough to be ejected so that the final part will conform to all of its specifications. Both the design of the part and the design of the mold are critical in insuring the right mold.

Both the design of the part and the design of the mold are critical in insuring the right mold. While the design of the part is not within the scope of this paper, it should nevertheless be thoroughly reviewed by those people directly involved in the design of the mold and in the molding operation [5].Before an injection mold is built the mold design must be established. Some basic considerations for the mold include the type of gating, the thermal control system, the type of ejection, the type of venting, and anticipated shrinkage. With these items in mind, the designer defines a parting line and adds the necessary features to the mold halves.When designing a mold for direct manufacture using SLS in the Sinterstation the cooling lines, ejector pin guides, gates, and runners can be included in the CAD design and built directly in the Sinterstation, see Figures 1 and 2.

Building these features directly into the mold saves time by reducing the man-hours spent preparing the mold for injection molding.Adding the cooling lines to the CAD file is an advantage because they can be placed near The alternative to crucial mold features. adding cooling lines to the CAD model is placing copper tubing in the back of the mold half prior to backfilling. The copper tubing can also be placed near critical mold features. tubing is more However, placing copper difficult the than placement of virtual cooling lines in the CAD model.andtimeconsumingEjector pin location plays a critical role in the success of the mold. A Copper Polyamide mold requires 30 to 40 percent more ejector pins than a standard aluminum tool. Add ing . more ejector pins makes part ejection easier which extends the lIfe of the mold. When applymg ejector pin holes in the CAD file, the holes should be made the size of the ejector pin. These holes are reamedFigure I.

When applymg ejector pin holes in the CAD file, the holes should be made the size of the ejector pin. These holes are reamedFigure I. Back surface of a mold core..452.out during the mold-finishing phase. After the shelling operation, extend the ejector pin guides be?"on.d the back surface of the mold to prevent the holes from becoming clogged when the backfill materIal IS added.Runner and gate designs are important to mold life and part quality. Gates for the Copper Polyamide molds should be larger than gates on an aluminum tool. The larger gate will keep the cavity pressure down. On large molds, fan gates are recommended. On smaller molds, edge gates are used.RunnerWhen the CAD design of the mold is complete, the mold half is shelled so that a backfill material can be added to aid in the conduction of heat away from the mold surface.

The shelling operation is performed either in the CAD package by operating on the solid like specialized model, Materialise's Magics RP [6] which operates on the In this step, the mold is hollowed out such STL file. that the external surfaces have a wall thickness of 3.8 mm. The bottom surface is then removed to provide access to the back surface of the mold face.orinaapplicationFigure 2. Working surface of a mold core.If the cooling lines are added to the CAD file, the wall where the cooling lines enter/exit should be 12.5 mm thick. is machinable the cooling line interface can be reamed and tapped for connecting the water source.Sincethe Copper PolyamideWhen using a low melting metal alloy as a backfill material, include a rib grid in the back of the mold to lock the alloy into the mold, see Figure 3. This rib grid is not necessary if a metal filled epoxy is used as the backfill material, see Figure 5.Figure 3.

This rib grid is not necessary if a metal filled epoxy is used as the backfill material, see Figure 5.Figure 3. Rib-grid to lock in backfill material.SLS ProcessingAfter the mold design is complete and a STL file is created, the part is built using the Selective Laser Sintering Process [7]. The Copper Polyamide material is processed with temperatures set points and scan parameters similar to those used to process DTM's other polyamide materials.The STL file in oriented so that the parting line is facing up with the longest dimension parallel to the x axis, then rotated 15 degrees around the x-axis. The purpose of angling the part is to add strength to small post features.Sacrifice geometry is added at the top and bottom of the build to control the cooling rate. The additional layer of scanned material acts as a heat fence reducing the thermal gradient within the part cake.

The additional layer of scanned material acts as a heat fence reducing the thermal gradient within the part cake. Without this barrier, the mold inserts would cool too quickly because of the good thermal conductivity of the Copper Polyamide material, resulting in curled or warped mold inserts. The sacrifice geometry can be any geometry that covers a majority of the build area and is at least 2 millimeters thick.453Mold finishingAfter the mold insert is built in the Sinterstation, it is finished and prepared for molding. The finishing steps include sealing the surface, finishing the surface, casting backfill material, machining ejector pins and cooling line connections, and aligning the insert in the mold base.The Copper Polyamide parts as built in the Sinterstation are porous. The inserts must be the molded plastic from sealed to prevent adhering to the surface and to prevent the conformal cooling lines from leaking.

The inserts must be the molded plastic from sealed to prevent adhering to the surface and to prevent the conformal cooling lines from leaking. A low viscosity epoxy or acrylate works well to seal the surface. It is important to select a sealant that can withstand high temperatures during molding. DTM recommends Imprex Superseal 95-1000A, which is a heat cure acrylate.the mold surface is After sealing the part, finished with a flexible sanding cloth, see Figure 4. Start with 220 grit, then 320 grit, and finally a 400 grit. Either wet or dry sanding techniques can be used. Apply another coat of L.... sealer to the mold surface after sanding.The backfill material recommended by DTM is Metspec 217. This material is a metal alloy with a melting temperature of 103°C. The Metspec is poured in layers to avoid softening and deforming the mold insert. When these steps are complete, the mold insert is inserted into the mold base. Different types of mold bases have been used successfully.

When these steps are complete, the mold insert is inserted into the mold base. Different types of mold bases have been used successfully. If using a two-plate type mold base, dowel holes can be added to the CAD model in the comers of the mold insert to aid in alignment. If using a one plate pocketed base, conventional methods are used to square up the insert before placing the mold insert into the pocketed base.Figure 5. Backfilling mold inserts with a metal filled epoxy.Molding TrialsThe mold inserts are fairly durable allowing a molding house to run the mold without special instructions. The durability of the Copper Polyamide mold inserts is similar to that of cast metal/epoxy tooling. The thermal conductivity of the material is good, 1.28 W/m-K at 40°C, and is also similar to that of cast metaVepoxy tooling.

The thermal conductivity of the material is good, 1.28 W/m-K at 40°C, and is also similar to that of cast metaVepoxy tooling. The good thermal conductivity, in combination with the conformal cooling lines and a metal backfill material, allows the mold to be run with normal cycle times, typically 25 to 40 seconds. The mechanical and thermal properties of SLS objects made with the Copper Polyamide are reported in Table 2.454_When molding, the mold should be at a consistent molding temperature before injection of the plastic. Try to get a full shot the first time. If there is a short shot, it is more likely to stick in the mold and cause damage. Mold release can be used at the beginning of the molding run. It is not necessary to apply mold release after every shot or late in the molding run. Several different release agents have been used with no problems.A number of common plastics have been successfully molded in the Copper Polyamide mold inserts.

Several different release agents have been used with no problems.A number of common plastics have been successfully molded in the Copper Polyamide mold inserts. The melting temperatures and processing temperatures for the plastics used in the mold inserts are listed below.

The melting temperatures and processing temperatures for the plastics used in the mold inserts are listed below. In most cases, the mold inserts showed little or no wear and can be used to mold additional parts.Table 1 Examples of plastics successfully molded in Copper Polyamide mold inserts.MaterialMelting Temperature, °CProcessing Temperature Range, °C (OF)Parts moldedPolyethyleneTm122 -137175 - 260 (350 - 500)50-350Polypropylene (PP)Tm160 -175190 - 285 (375 - 550)50-20010-40% talc filled PPTm158 -168175 - 285 (350 - 550)50-35040%GF PolypropyleneTm168230 - 285 (450 - 550)30 -150ASS (medium impact)Tg102 -115200 - 275 (390 - 525)40 -150ASS/PVCTg175 - 205185 - 210 (370 - 410)100 - 200Case Study #1: Brake fluid reservoir molded in polypropylene [8]Last spring at BASTECH, Inc [9], an automotive OEM client offered a challenging assignment.

It needed 20 to 50 sets of brake reservoir parts produced in the intended production material, polypropylene. The parts had to withstand prolonged contact with brake fluid. Moreover, they had to perform like production parts during functional testing on a prototype vehicle.BASTECH had faced a similar situation a year ago. Back then, the company had used 3D Systems' AIM Tool process to create prototype tooling. The results were only marginally successful. After 10 shots with the AIM Tool they began to see deformation of the parting line and the mold itself. The run produced only 40 polypropylene parts, with each set becoming slightly more deformed.Wanting something better for this molding run, BASTECH used the Copper Polyamide material. The brake reservoir consisted of two pieces, which eventually would be welded together. Two tools were required to make these parts. One tool was made up of 13 pieces; the other was made up of 6 pieces.

Two tools were required to make these parts. One tool was made up of 13 pieces; the other was made up of 6 pieces. Eight pieces of the thirteen-piece tool were made using the Copper Polyamide (Figure 6), as were five pieces of the six-piece tool. The remaining pieces, very simple inserts or sleeves, were machined out of metal. The final welding, measured approximately 110 mm by 154 mm by 41 mm.part,afterFigure 6. STL files used to build the mold halves and inserts.455BASTECH ran 75 sets of parts on the Copper Polyamide tools and noticed very little wear. They started to see a little degradation in a few areas, but could easily have gotten 100 to 150 sets of parts from the molds. The SLS Copper Polyamide tooling withstood an injection molding pressure of 400 psi and injection molding temperatures of 230°C. They stated that the molds were very sturdy which impressed them the most. Overall, both BASTECH and there customer were very pleased.

They stated that the molds were very sturdy which impressed them the most. Overall, both BASTECH and there customer were very pleased. The molded parts are pictured in Figure 7 next to one of the mold halves.The SLS Copper Polyamide tooling well It took an extra outlasted the AIM tooling. day to fit, finish, and polish the SLS tooling; but the resulting tools were more stable, not deforming during molding, and the parts produced looked much better. The time and costs for producing tools using Copper Polyamide in the SLS process or Epoxy in the SLA process are roughly the same.Comparing the timing and costs of an SLS a Copper Polyamide conventional steel or machined production tool offers more compelling numbers. The SLS molds can be produced at roughly one fourth the cost in only one half the time when compared to steel tooling. These savings are important when only 50 to 500 molded parts are needed.tooltothatofFigure 7.

These savings are important when only 50 to 500 molded parts are needed.tooltothatofFigure 7. Parts molded by BASTECH in the Copper Polyamide molds pictured next to one of the mold inserts.Case Study #2: Glass guide molded in nylon 6,6 [8]The Rover Group [10], at their Rapid Prototyping and Tooling Department situated on the campus of the University of Warwick, has participated as a beta test site for the Copper Polyamide material. Rover has studied, among other things, how tooling made via the SLS process with the new Copper Polyamide material compares to alternative low cost tooling routes.One of the parts selected for Rover's comparison was a glass guide (approximately 90 mm x 60 mm x 25 mm). This part is used to ensure the automobile door window locates in the rubber seal when closing. The intended production material was nylon 6,6.

This part is used to ensure the automobile door window locates in the rubber seal when closing. The intended production material was nylon 6,6. They had already created the tool geometry in cast resin, so when they created a Copper Polyamide tool with the SLS process it was purely for the sake of comparison.Rover designed the single cavity mold inserts as solid objects on a CATIA workstation. The STL files were then shelled to a wall thickness of 2.5-mm using Materialise Magic RP software. The inserts were produced in one of Rover's two in-house Sinterstation 2500 systems. Once removed from the Sinterstation, the hollow tool inserts were backed with an epoxy resin loaded with aluminum powder and granules. The tool was subsequently machined to fit one of Rover's standard injection molding bolster sets.To trial the Copper Polyamide mold inserts and to ensure the tool was working properly, Rover shot an initial run of 33 polypropylene parts.

Once confident that the tool was working properly, they shot another456117 nylon 6,6 parts into the Copper Polyamide tool. Nylon 6,6 parts molded in the Copper Polyamide tool are pictured in Figure 8.Rover was pleasantly surprised with the results. They had not expected the Copper Polyamide tooling to withstand the high injection temperature of nylon 6,6 which is Some minor degradation of the 285°C. Copper Polyamide tool began to occur after 20 shots. The degradation was most noticeable around the sprue puller where the sharp edges began to erode. This area was directly across from the injection sprue where the mold reaches the highest temperatures. Dr. Illston at Rover stated that the degradation of the tool part exaggerated because was geometry was an early prototype design. The glass guide part featured a section of approximately 1 cm3 that retained considerable heat during injection molding.

The glass guide part featured a section of approximately 1 cm3 that retained considerable heat during injection molding. Increased localized heat build up resulted in additional erosion of the male part geometry where the small radii were intended.alsothePolyamide tool. 117 parts molded.The final result ofthis project was that 60 nylon glass guides made from the Copper Polyamide tools have been installed on Rover Group's prototype vehicles. Under the circumstances, the performance of the tool was far better than expected.Rover has continued to explore the advantages of the SLS-generated Copper Polyamide tooling. Another project resulted in 300 parts molded with 30% talc filled polypropylene. The part is a speaker fret (or speaker cover) which measures 38 mm in diameter and features a complex geometry, see Figure 9.

The part is a speaker fret (or speaker cover) which measures 38 mm in diameter and features a complex geometry, see Figure 9. During the production of these parts, the tool showed no wear or part The injection temperature of the polypropylene material was 240°C.flashing.Rover has worked with a number of rapid tooling methods. the Copper Polyamide mold inserts compare more favorably to cast resin tools. Although the performance of the Copper Polyamide tools is comparable to cast tools because the thermal conductivity of the two material systems are similar, there are advantages to building the mold inserts directly in the Sinterstation from the CAD data. When asked which method is faster, Dr. Illston says it all depends on the data available. If the CAD data is available for the mold inserts then the Copper Polyamide route is quicker. If the CAD data is only available for the part, then it is quicker to create a RP master pattern and cast the resin tool.

If the CAD data is only available for the part, then it is quicker to create a RP master pattern and cast the resin tool. Rover is currently evaluating the cost and time comparisons between the two processes.Their view is thatFigure 9. Speaker fret parts molded with 30% talc filled polypropylene. 300 parts molded in the Copper PA tool.457ConclusionThe Copper Polyamide material and the Sinterstation system can rapidly produce limited life prototype mold inserts. Prototype quantities of production-quality molded plastic parts can be manufactured with these metal-based mold inserts.The important attributes of a short run tool are durability, molding cycle times, and speed. The durability of the Copper Polyamide tools is comparable to epoxy/metal tooling and is much better than Direct AIM tools. The molding cycle times for Copper PA are similar to metal tooling. This is important because cycle times can affect the mechanical properties of the molded parts.

The molding cycle times for Copper PA are similar to metal tooling. This is important because cycle times can affect the mechanical properties of the molded parts. Mold inserts are created relatively quickly with the Copper Polyamide material. Molds can be ready for molding in 5 workdays; this includes all processing steps required to move from a STL file to molded parts.Table 2. Copper Polyamide product specificationUNITSASTMTEST METHODBUILT in SinterstationPHYSICAL PROPERTIES@15°angle!FLAT;ityg/cm30-7923.45THERMAL PROPERTIES..." " PIV..kL.9A? M.P.9.".(§§..p~JL.." . " ".." "PIV...~LJ..:.§?"MP.9.j?..§.4...P.§D."... "... ..........."".Ib~.r.r.D..i:lJ"G.9..Q.g.~~!.i.y.i.ty. ..."........... """.$p~~i.fi~."H.~9.L .."........." ".G.9.~m~i.~.Q.t..I~.E?r..'!.1.9.L~~p.i:l.Q.~.i.9.Q40° to 150°C40° to 150°C30° to 150°C"."." ~!.~~.~.9.." m/mrC".".0-648 ....................··..······....···1..·············0-648. "",,1.........

RECT LASER DEPOSITION OF Ti-6Al-4V FROM ELEMENTAL POWDER BLENDSLei Yan1, Xueyang Chen1, Wei Li1, Frank Liou1, and Joe Newkirk21Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409 2Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractA thin-wall structure composed of Ti-6Al-4V has been deposited using direct laser deposition (DLD) from blended Ti, Al, and V elemental powders. The microstructure and composition distribution along the build height direction were intensively investigated using optical microscopy, scanning electron microscopy/energy dispersive X-ray spectroscopy (SEM/EDS), and Vickers hardness testing. The microstructures of the as-deposited Ti-Al-V were studied using EDS to determine appropriate weight percentage for Al and V in the blended powders before mixing.

The microstructures of the as-deposited Ti-Al-V were studied using EDS to determine appropriate weight percentage for Al and V in the blended powders before mixing. The effects of laser power and laser transverse speed on the microstructure were investigated and optimized laser processing parameters were concluded.1. IntroductionDirect laser deposition (DLD) is a viable and promising technology that has revolutionized manufacturing complex and hard-manufactured material components. DLD manufactured parts are precisely built by adding material layer by layer, usually in powder form. Conventionally, a prealloyed powder, in which each individual powder particle has the composition of the desired alloy composition in the final part, is used in the DLD technique [1]. A different or an improved approach of laser deposition is using mixed elemental powders to flexibly create materials in situ with the development of powder feedstock.

A different or an improved approach of laser deposition is using mixed elemental powders to flexibly create materials in situ with the development of powder feedstock. With the blends of elemental powders, each individual powder particle has an elemental composition of an element present in the desired final alloy composition. Upon mixing, the sum of all powder particles gives the desired alloy composition in the final part [1]. A variety of materials including metals, alloys, and ceramics have been made with elemental powders such as SS316, titanium-based alloys, nickel-based superalloys Inconel 625, and Fe-Cr-Ni, Fe-Co-Al system alloys [2-8]. Those successes could potentially reduce the costs of processing complex components from powder preparation post-processing to a large extent. In addition, using elemental powder could also benefit the deposition of functionally graded materials (FGM), such as from metal to metal, Ti-V (Mo), and metal-ceramic, Ti-TiC [9-10].

In addition, using elemental powder could also benefit the deposition of functionally graded materials (FGM), such as from metal to metal, Ti-V (Mo), and metal-ceramic, Ti-TiC [9-10]. However, only limited studies on Ti-6Al-4V by elemental powder with the DLD method were reported. Present work is focused on processing parameters of laser transverse speed, laser power, and their effects on microstructure, Vickers hardness, and material composition distribution along build height direction.2. Experimental proceduresThin-wall samples were fabricated at MST laser-aided manufacturing process (LAMP) lab using the DLD system, which consists of an argon-purged chamber, 1 kW Nd-YAG fiber laser,side nozzle powder feeder, and 3-axis numerical control work table. Fig.

Fig. 1 shows the experimental set-up of the MST-LAMP DLD system.Commercially pure Ti, Al, and V powders were weighed out and two sets of powder blends were prepared according to weight percentage ratios of 90:6:4 and 84:8:8 after being dried in the furnace for 30 minutes at 200°C. Two sets of blends were mixed in a sealed bottle using a Turbula® mixer for 10 minutes. The characteristics of the powders are listed in Table 1.The thin-wall samples were deposited on 2 inches long, ½ inch wide, and ¼ inch thick Ti- 6Al-4V substrates. Based on preliminary experiments, laser power from 550 to 750 W, laser transverse speed from 200 to 400 mm/min, the same powder feed rate, and laser beam diameter 3 mm were chosen for the experiments. The experimental design and processing parameters are listed in Table 2.Cross section of the deposits and substrates were EDM cut, mounted, and then polished according to standard metallographic preparation of Ti-6Al-4V.

The samples were etched by Kroll’s reagent consisting of 92 ml distilled H2O, 6 ml HNO3, and 2 ml HF to reveal α and β phases for further analysis.Fig. 1: Schematic of the MST-LAMP DLD system.Table 1: Powder charactersPowder Ti Al VParticle AppearanceirregularParticle Size/μm 106 165 175Sample No.1 2 3 4 5Table 2: Experimental design and DLD processing parameters PowerScanning Speed (mm/min) 200 400 300 200 200LayerLayer Thickness (mm) 0.2 0.2 0.2 0.2 0.2750 650 650 650 55020 20 20 20 20Vickers hardness was first tested on a Vickers hardness tester with 1 Kg load and 10 seconds dwell time. Vickers hardness sampling points were distributed along three zigzag routes at 0.3 mm intervals and the Vickers hardness at certain positions was taken from the average VHN per each row with the same height. Fig.

Fig. 2 represents how the Vickers hardness tests were conducted.The microstructure was studied under the optical microscope (OM) to determine microstructure types and the variation of grain size. The composition distribution along the build height direction was investigated by employing SEM/EDS with Helios 600. Both line scan and area scan were used for the EDS test. For line scan test, lines in the middle of the deposits along the build height direction with length of 160 μm were drawn to detect the homogeneity of material distribution. During area scan tests, 200 μm by 200 μm square areas with center point intervals 300 μm and 500 μm were scanned to determine Ti, Al, and V composition distribution in the as- deposited alloy from top to bottom. The two different intervals 300 μm and 500 μm are for deposits made from powder blends with weight percentage ratio 90:6:4 and 84:8:8 separately. A smaller interval of 300 μm was chosen for short deposit height of the tested samples.

A smaller interval of 300 μm was chosen for short deposit height of the tested samples. A Ti-6Al-4V bulk containing 6.33 Al, 4.10 V, 0.17 Fe, 0.19 O, 0.011 C, 0.013 Cr, 0.012 Ni and 0.02 Si and balanced Ti (all composition in Wt.%) was used as the standard for both line scan and area scan during the EDS tests, which made test accuracy within ±1% [11].Fig. 2: Vickers hardness test pattern.3. Results and discussion 3.1 MicrostructureFig. 3 shows the microstructure of DLD Ti-6Al-4V from blended elemental powders with laser transverse speeds 200, 300, and 400 mm/min at a given laser power of 650 W. The microstructure contains primary α, secondary α and β. The microstructure varies from bottom to top of the deposits, but a general trend showed the microstructure would be finer closer to the bottom.

The microstructure varies from bottom to top of the deposits, but a general trend showed the microstructure would be finer closer to the bottom. A steady state was reached from middle to top where not too much difference could be found in the microstructure and basically shows Widmanstätten basketweave, which tends to form with an increasing cooling rate from the β-phase field [12-13].abcdeFig. 3: Microstructure of deposits at laser transvers speed: (a) 200 mm/min, (b) 300 mm/min, (c) 400 mm/min, (d) bottom of deposit at 300 mm/min, (e) top of deposit at 300 mm/min.The microstructure variation for different laser transverse speed can be explained by the introduced volumetric energy density, which has a high influence on microstructure and its relationwith laser power P (W), laser transverse speed v (μm/s), hatch spacing h (μm) and layer thickness d (μm). This can be expressed as: [14].

This can be expressed as: [14]. During the experiments, hatch spacing and layer thickness were fixed so increasing laser power and/or decreasing laser transverse speed would increase energy density and leads to a higher cooling rate, which would benefit finer grain formation.E P v d h / ( * * )Fig. 3 (a) to (c) shows the basketweave microstructure chosen from the steady state section of three deposits with different laser transverse speed, which reveals that the grain size increased with the laser scanning speed. From the definition of energy density, the higher the laser transverse speed, the lower the cooling rate, which leads to a coarser microstructure [15].Fig. 3 (d) and (e) shows the microstructure at the bottom and top of the deposit with 300 mm/min laser transverse speed, which reveals that at the lower end of the deposit near the substrate, the grain size was smaller than the upper side.

The same trend described above happened at scanning speed 200 mm/min and 400 mm/min, too. The reason was the closer to the substrate, the higher cooling rate happened, and contributed to the formation of finer grains.Similar phenomenon for laser power 550, 650, and 750 W at fixed laser transverse speed 200 mm/min was observed and the difference between microstructures from different deposits or different positions of the same deposit all could be explained based on the energy density and cooling rate. Therefore, it was concluded that higher laser power and/or slower laser transverse speed tends to benefit the formation of finer microstructures.3.2 Vickers HardnessThe average Vickers hardness values of as-deposited Ti-6Al-4V from blended elemental powders are shown in Fig.

4 to represent Vickers hardness with respect to position in the deposits and the relationship with variations of laser power and laser transverse speed.Ti-6Al-4V substrate Vickers hardness was first taken near the bottom edge that is away from heat affected zone (HAZ) and the Vickers hardness ranges 346.47±9.96 VHN. It was observed that on the substrate side for each laser transverse speed and laser power, the Vickers hardness values increased when it went into the HAZ. When it goes into the deposit zone, an obvious relationship for both laser transverse speed and laser power with Vickers hardness was observed. Fig. 4 (a) shows Vickers hardness of the deposit at laser power 550 and 650 W doesn’t vary too much and almost stayed constant, Vickers hardness varied 449.64±7.18 and 454.20±7.10 VHN, separately.

Fig. 4 (a) shows Vickers hardness of the deposit at laser power 550 and 650 W doesn’t vary too much and almost stayed constant, Vickers hardness varied 449.64±7.18 and 454.20±7.10 VHN, separately. For laser power at 750 W, Vickers hardness ranged 506.04±22.52 VHN and showed an increase from the first several layers, then stayed stable from the middle to the top of the deposit. This is as expected because the higher the laser power, the higher the energy density and higher cooling rate. The higher the cooling rate, the finer the microstructure and which leads to higher Vickers hardness. Fig. 4 (b) reveals the relationship between the laser transverse speed and Vickers hardness. Vickers hardness ranged 459.04±9.16, 454.18±8.76, and 444.86±16.74 VHN for laser transverse speed 200, 300, and 400 mm/min, respectively. Vickers hardness of the three deposits stayed stable at three different laser transverse speeds.

Vickers hardness of the three deposits stayed stable at three different laser transverse speeds. It was clear that the decrease in laser transverse speed led to the increase of Vickers hardness and that it also influenced the cooling rate. As the laser transverse speed decreased, energy density and the cooling rate increased. Thus, the lower laser transverse speed was the direct reason for fine grains and higher Vickers hardness.abFig. 4: (a) Vickers hardness vs. laser power, (b) Vickers hardness vs. laser transverse speed.3.3 EDS TestTi, Al, and V powders were first mixed with the designated weight percentage ratio of 90:6:4 to get target material Ti-6Al-4V. Area scan and line scan were employed to check the composition distribution along the build height direction. Fig. 5 (a) shows from substrate to top of the deposit Al and V weight percentage decreased and in the opposite direction Ti increased.

Fig. 5 (a) shows from substrate to top of the deposit Al and V weight percentage decreased and in the opposite direction Ti increased. Al and V curves were zoomed in and show their weight percentage variation against position in Fig. 5 (b). Two marked zones in Fig. 5 (b) mean industry qualified Al and V weight percentage for Ti- 6Al-4V, which is Al ranges from 5.5-6.75% and V 3.5-4.5%. At deposit height from 0-0.9 mm only two areas that have the qualified composition weight percentage; this may have been caused by Al and V diffusion from the substrate when the melting pool was generated. Above height 0.9 mm, Al and V weight percentage shows the trend that the values stayed stable around Al 5% and V 2%, which means the as-deposit alloy is actually Ti-5Al-2V and not target material Ti-6Al-4V. Line scan was also employed during the EDS test to define composition homogeneity along the build height direction. The line scan result in Fig.

Line scan was also employed during the EDS test to define composition homogeneity along the build height direction. The line scan result in Fig. 5 (c) shows a homogenous chemistry throughout the deposit, but the number of counts doesn’t mean an exact measurement of chemistry; the line scan is just a qualitative analysis, but not a quantitative analysis. Improvements could be made by increasing Al and V weight percentage of the blended powders or generating a much bigger melting pool to increase the chance for Al and V powders to fall into.In order to increase Al and V weight percentage in the deposit, a trial and error method was used and a new weight percentage ratio of 84:8:8 for Ti, Al, and V powers was mixed. Fig. 6 shows the EDS results of the deposit based on the new designed weight percentage ratio. Fig. 6 (a) shows the weight percentage of Al and V slightly increased from the substrate to about 1 mm height and then stayed almost unchanged. Fig.

Fig. 6 (a) shows the weight percentage of Al and V slightly increased from the substrate to about 1 mm height and then stayed almost unchanged. Fig. 6 (b) is the zoomed in curves of Al and V and shows that the first 1 mm height of the deposit has qualified Al and V weight percentage and also may diffuse from the substrate via the melting pool. Above 1 mm height, Al and V weight percentage fell into narrow zones 6.75-7% (6.9±0.21%) and 4.5-5% (4.72±0.10%), which means the 84:8:8 weight percentage ratio could make Ti- 6.9Al-4.7V and get pretty close to the maximum limit of industry qualified Ti-6Al-4V.It could be assumed that a little decrease of Al and V content ratio in the mixes from 84:8:8 ratio, such as to 86.5:7:6.5, could get qualified Ti-6Al-4V deposits. The reason for the lack of Al in the deposits may be caused by Al powder’s high reflection ratio of the laser beam and lighterweight compared to Ti and V powders.

The reason for the lack of Al in the deposits may be caused by Al powder’s high reflection ratio of the laser beam and lighterweight compared to Ti and V powders. The lighter weight may lead to Al powder flowing and falling outside the melting pool. Low V weight percentage means less V powder was falling into the melting pool during the deposition process and this may come from the characters of V powder being used during the experiment [16]. V powder’s largest particle size and density would cause some of it to fall behind or outside the melting pool and result in the lack of element V. This is a common phenomenon for blended elemental powders, especially when different elemental powders’ average particle size and density has significant variation.

This is a common phenomenon for blended elemental powders, especially when different elemental powders’ average particle size and density has significant variation. DLD would benefit from multiple coaxial powder nozzles in this situation [8], which means each nozzle could be set to be perfectly suitable for specific elemental powders to fall into the melting pool.abcTitanium Kα1Vanadium Kα1 (upper line)Aluminum Kα1 (lower line)Fig. 5: EDS results for Ti-6Al-4V: (a) Composition distribution along build height direction, (b) zoom in of Al and V curves in Fig. 5 (a), (c) Line scan along build height direction.abFig. 6: EDS results for Ti-8Al-8V: (a) Composition distribution along build height direction, (b) zoom in of Al and V curves in Fig. 6 (a).4. ConclusionsThe influence of the laser transverse speed and laser power on microstructure, Vickers hardness, and composition distribution of Ti-6Al-4V made from elemental powder by DLD process were studied.

Increasing laser power or decreasing laser transverse speed tended to benefit the formation of finer grains. Vickers hardness is related to grain size, and the finer the grain size, the higher Vickers hardness the material reveals. Thus, increasing laser power or decreasing laser transverse speed increased Vickers hardness. The EDS tests results show that increasing Al and V weight percentage from 6% and 4% of the powder blends to 8% and 8% would increase Al and V weight percentage in the deposit and make it much closer to the industry qualified Ti-6Al-4V.AcknowledgementsThis work was conducted as part of NASA’s project through EPSCoR Grant#NNX13AM99A. The support from Boeing is also gratefully acknowledged.References [1] Clayton, Rodney Michael. The use of elemental powder mixes in laser-based additive manufacturing. 2013. [2] Brown, Clyde O., Edward M. Breinan, and Bernard H. Kear. "Method for fabricating articles by sequential layer deposition." U.S. Patent No. 4,323,756.

Direct laser fabrication of this component fiom a bed a loose powder results in significant cost savings and improved performance over the currently employed production technique. The technology has been demonstrated by fabricating a prototype lot of 100 blade tips, which will be subjected to an engine test. This is the first instance of a direct fabrication method applied to the production of functional engine hardware. This research was funded by the United States Air Force contract F33615-94- C-2424 titled "Affordable Turbine Blade Tips".SUMMARYThis program has contributed significantly towards meeting the Advanced Technology affordability goals for turbine engine acquisition. These goals require engines to be developed and maintained at a reduced cost. To meet some of these cost reduction goals, advanced manufacturing methods such as solid eeeform fabrication (SFF) must be applied.

To meet some of these cost reduction goals, advanced manufacturing methods such as solid eeeform fabrication (SFF) must be applied. Although only one small turbine engine component is being discussed here, SFF in general has great potential for impacting cost reduction goals. The work performed is developmental, whch means fewer parts are produced regardless of the component. Since SFF costs are nearly independent of the number of ~ a r t s ' , it will allow components to be fabricated at reduced cost. Since SFF has the potential to significantly reduce time from design conception to fabrication, it will also allow more design iterations to take place.The use of SFF to directly produce fbnctional components eliminates the costs and time associated with raw material acquisition, conversion and manufacture.

The final goal of developing such a process would be to have the ability to control the microstructure with respect to precipitate size, density, porosity size and distribution, grain size, crystallographc texture and solidification modes. Such process control would allow production of components tailored with functionally graded microstructures to enhance properties and performance in specific regions ofthe component. This would represent the pinnacle of life and performance enhancement, and cost and schedule reduction.SFF can also be applied to tooling production applications such as investment castings. Investment castings are used for complex geometries such as airfoils, where SFF can provide a means for prototyping tooling or patterns. These prototypes assist in production trials and iterations that are necessary to finalize geometry and processing2.

These prototypes assist in production trials and iterations that are necessary to finalize geometry and processing2. Whether it is used for component or tooling manufacture, SFF is a critical area and a technology that is required to reduce the cost and lead time for producing turbine engine components.INTRODUCTHONUntil recently, no work was reported on direct selective laser sintering (SLS) of high performance materials such as Nickel and Cobalt base superalloys, superalloy cermets9, Titanium base alloys3 and monolithic high temperature metals such as ~olybdenum~. These materials are used for high performance components that typically experience high operating temperatures, high stresses and severe oxidizing or corrosive environments. Direct SLS, with its ability to produce components in such materials is especially useful for functional prototype, low volume or "one of a kind" production runs.

Direct SLS, with its ability to produce components in such materials is especially useful for functional prototype, low volume or "one of a kind" production runs. To manufacture a typical prototype lot of 100 superalloy cermet abrasive turbine blade tips, direct SLS was shown9 to achieve acceptable microstructure and properties with 80% cost savings over the traditional method. Automotive and aerospace industries face typical lead times of several weeks for functional, metallurgical quality prototypes. Direct SLS fabrication can lower cost and drastically reduce lead times by eliminating pre- processing and post-processing steps, and by eliminating the need for specialized tooling.Since its inception in the 1 930ts, the gas turbine engine has grown to be the workhorse power plant of modem aircraft. Advanced technology propulsion system are very important fi-om an overall system affordability viewpoint since higher engine performance can result in a smaller airfi-ame.

Advanced technology propulsion system are very important fi-om an overall system affordability viewpoint since higher engine performance can result in a smaller airfi-ame. SFF technologies are advanced manufacturing methods that can significantly impact cost reduction of engine components. SFF is striving towards decoupling cost from volume5 by making it possible to produce the first unit at a recurring cost equal to the hundredth unit6. To demonstrate SFF's cost-effectiveness, a superalloy cermet abrasive blade tip ( m T ) was chosen for direct fabrication technology development and demonstration.. . .ing leakage of the gas away from and around the working gas path. Therefore, clearance between rotating and static parts is critical. This clearance changes with component expansion and contraction due to the thermal cycling experienced in gas turbines.

Therefore, clearance between rotating and static parts is critical. This clearance changes with component expansion and contraction due to the thermal cycling experienced in gas turbines. One of the primary methods developed to compensate for this expansion and contraction is an abradable seal This seal system works by a t t a c h g an ABT to the tips of the turbine blades. The stationary turbine shroud is coated with a abradable ceramic, which is abraded by the abrasive blade tip during engine operation to form a gas path seal around the rotating component. The abradable ceramic liner is relatively inexpensive and simple to manufacture. However, the abrasive cermet (0.9- 1.5 mm thickness) is difficult to manufacture on a production scale and contains expensive components. The cost of the rawGas turbine efficiency is highly dependent upon2materials makes scrap highly undesirable. Due to the difficulties encountered9, a more reliable method for producing this material is desired.

Due to the difficulties encountered9, a more reliable method for producing this material is desired. A direct SLS process for producing the cemet abrasive can provide a uniform, repeatable manufacturing process that would eliminate much of the manual labor, reduce scrap and make all unsintered surplus materials reusable.OBJECTWESPrior research9 in direct laser processing ABT cermet material had shown the capability of directly laser processing cermet material to obtain microstructure and properties that were consistent with production specifications. The ultimate goal of ongoing research was to complete technology development and demonstration of a direct laser manufacturing process capable of repeatably producing the ABT component to dimensional specifications, with microstructure and properties equivalent or superior to current production material.

A prototype lot of 100 turbine blade tips would be manufactured to demonstrate the developed process and the associated efficiency.During process development, the objectives were fourfold. The first objective was to investigate and establish a variety of material composition variants and processing conditions, and their effects on solidification microstructure, and mechanical properties. The second objective pertained to identification and specification of optimal processing windows for manufacturing the ABT component in a single pass laser scan. Under this objective, the optimal processing conditions would result in filly dense, fully melted and rapidly resolidified material free of solidification defects such as hot tearing, pinholing, entrapped porosity, grit floating or segregation. A third objective was to demonstrate reusability of unprocessed material.

A third objective was to demonstrate reusability of unprocessed material. Atta* ' this objective would direcly impact cost effectiveness of the laser based manufacturing proceing scrap of expensive material constituents. The fourth and final objective was to prove consistency and repeatability of the laser process by producing a large number of ABT components in several experimental runs under identical processing conditions.MATERULS AND METHODSCermet SpecificationsThe cermet in production is vacuum sintered using a Silicon and Boron containing Cobalt based braze alloy (Amdry 788) and subsequently bonded to the blade using a similar braze. tion of the low melting point braze constituent could extend oxidation life and possibly allow higher operating temperatures. Higher operating temperatures could be achieved by replacing the braze used for bonding the ABT to the blade by a higher temperature diffusion braze.

Higher operating temperatures could be achieved by replacing the braze used for bonding the ABT to the blade by a higher temperature diffusion braze. Currently, interactions between the component braze and the bonding braze prevent the use of a diffusion braze alloy. The goal of developing a brazeless cermet composition for direct laser processing was achieved during previous work9. This cermet is composed of Mar-M-247 nickel superalloy matrix (Table I), GE Borazon 510 (titanium coated cubic boron nitride) and Alundum 90 (titanium coated alumina). The thickness and chemical quality of the titanium coating on the abrasive grit is critical. Poor chemical quality of the coating causes contamination of the molten matrix alloy either through dissolution alone or in combination with spallation.

Poor chemical quality of the coating causes contamination of the molten matrix alloy either through dissolution alone or in combination with spallation. Atitanium coating of non-uniform or insufficient thickness results in poor metallurgical bonding with the nickel matrix, and grit floating out of the melt due to large density and surface energy diRerences.To meet homo geneity requirements, the sint er ed material must exInibit no discrete layering or agglomeration of abrasive particles which would result in grit pullout and associated decrease in seal efficiency. Porosity requirements allow for no more than 20% linear porosity as determined from a metallographic montage using a line intercept method. If the material meets these specifications, coupons are removed and brazed into the gauge section of a CMSX-4 single crystal test bar conforming to ASTM-E 139 and subjected to an application specific stress-rupture test.Table 1.

Nominal composition of matrix material used in the cermet.Process Apparatus and MethodsA laser processing machine dedicated to the production of ABTs was designed jointly by the University of Texas and Allison Engine Company. The principle of operation of this machine is similar to a generic SLS machine. However, it has a few distinguishing features. Since one objective of ABT process development was to produce the full thickness (0.9-1.5 m) in a single laser pass, this machine does not provide multiple layer powder delivery. This machine is equipped with a 250 Watt CW Nd:YAG laser, high temperature powder heating capability and controlled atmosphere. Metal powders can be heated up to 1000°C in this workstation. This workstation has the capability of maintaining high vacuum (< 10'~ Torr) as well as high purity atmospheres of select gases.

Metal powders can be heated up to 1000°C in this workstation. This workstation has the capability of maintaining high vacuum (< 10'~ Torr) as well as high purity atmospheres of select gases. Custom software for CAD geometry processing was used to convert the electronic drawing of the demonstration ABT component into the desired laser scanning sequence. The machine control software allows the user to control and set processing parameters including preheat temperature cycle and setpoint, laser power, scan speed and scan spacing.Material preparation for the direct SLS method consists of combining by weight 73% superalloy, 18% Ti coated alumina, and 9 % CBN. This mixture is blended for 4-6 hours prior to loading in the SLS chamber. The blended mixture is poured into a 4" diameter, 0.060-0.100" deep processing dish and leveled to produce a uniform thickness powder layer.

The blended mixture is poured into a 4" diameter, 0.060-0.100" deep processing dish and leveled to produce a uniform thickness powder layer. The powder mixture was then baked out under high vacuum (< 5 xloW5 Torr) and with 25" C to 875" C temperature ramp over 30 minutes prior to SLS processing.Direct SLS involves directly melting and consolidating selected regions of a powder bed to form a desired shape having full density. This method of fabrication involves melting the component matrix and obtaining the appropriate amount of flow f?om the molten material. The appropriate amount of flow is critical. It can be described as the produces a highly dense part maintaining dimensional tolerances an s other defects such as hot tearing, pitting, plnholing and solidification cracking. A number of experiments wereconducted to evaluate the influence of the most important SLS process parameters on the solidification behavior of this material composition.

A number of experiments wereconducted to evaluate the influence of the most important SLS process parameters on the solidification behavior of this material composition. These parameters are atmosphere, preheat temperature, scanning sequence and energy density. The appropriate amount of flow is controlled by several factors such as atmosphere, degree of preheat and three characteristic variables affecting laser energy density: laser power, scan spacing and scan speed.Rectangular coupons 1" long X 0.5" wide were produced by raster scanning across the width. These samples were used for process parameter development and mechanical testing. Tip shape processing was done by two alternate methods. The first method involved raster scanning the shape of a tip on a powder bed container. The second method was to laser process a rail shaped tip out of powder contained in a mold.

The first method involved raster scanning the shape of a tip on a powder bed container. The second method was to laser process a rail shaped tip out of powder contained in a mold. The mold material directly affects the thermal profile generated and thus can be used in conjunction with preheat and laser scanning preheat to assist in control of the solidification microstructure of the component. The physical properties of the powder container influence the microstructure of the part dramatically. A ceramic powder container insulates the bottom surface of the powder bed and sets up a fairly unifonn thermal gradient which results in partial directional solidification or epitaxial growth. A more conductive powder, e.g.

A more conductive powder, e.g. a steel or superalloy container can produce a very fine dendrite spacing in the solidification microstructure,The laser scanning sequence significantly affects the thermal profile of the part during processing, therefore it will have a significant affect on tendency of the material to tear, crack and separate. The scanning sequence incorporated on actual parts will be dependent upon part geometry and modified in an attempt to provide a uniform thermal profile. This experiment involved only two variables on rectangular coupons, rastering the beam on continuously back and forth, and intermittently shutting the laser off and only rastering left to right. Scanning sequence can dramatically affect local solidification t b e , which affects dendrite spacing or solidification mode.Traditional metallurgy practices dictate that the superalloy cermet material requires processing in a high vacuum atmosphere.

However, a quick evaluation of the affects of different atmospheres was performed to document the degree of degradation. This consisted of perforrning trials in different types of atmospheres that included rough vacuum (of order 30mTorr), rough vacuum purged with argon, high vacuum (< 5 ~ 1 0 - ~ Torr) and high vacuum followed by argon-2% hydrogen backfill.Experiments were conducted to determine the affects of preheating the powder bed in the range of 3 50°C to 1000°C vs. no preheat. The experimental results provided two fmdings. First, preheat was beneficial in that it hastened outgassing of the powder. Second, if a well outgassed powder was placed in the powder bed, the preheat improved the surface characteristics, wetting and flow, and minimized hot tearing or cracking. In general the components manufactured were more consistent in surface texture quality if a preheat is applied.

In general the components manufactured were more consistent in surface texture quality if a preheat is applied. Preheat could also be used to manipulate the thermal gradient during solidification which affects dendrite spacing and the mode of solidification.The life of the ceramic mold was not as long as desired, therefore a metallic mold was used to produce the components. A partial production lot of ABTs produced via this method is shown in Figure 2. The production yield was excellent, in excess of 95%. The mold was 0.090" deep and yielded parts that were in the 0.077 to 0.084" thck range. Subsequently, one surface was ground flat for bonding. The tips attained a final thickness in the 0.073 to 0.077" range.Process Parameters for Production.Precipitate strengthened nickel base superalloys are considered difficult to weld or work with in the molten state due to their tendency to strain age crack10911.

Therefore, SLS processing parameters should be aimed at atta the most uniform thermal profile possible during processing to e the tendency for hot tearing and solidification cracking due to residual stresses and stresses resulting from aging during processing. This implies that a very fine scan spacing with a high scan speed should be incorporated. These two parameters also affect the surface roughness of the sintered coupon. A fine scan spacing will provide a relatively uniform smooth surface. Scan speed effects on surface roughness are dependent upon the overall energy density and the associated residence time in the molten region. For a given material a high scan speed with a high energy density which would produce extensive superheat or a long molten residence time will produce a poor surface finish. The same scan speed with a lower energy density and thus a lower degree of superheat will produce a better surface finish.

The same scan speed with a lower energy density and thus a lower degree of superheat will produce a better surface finish. In general excessive superheat will allow growth of morphological instabilities in solidification fiont s that produce non-uniform surface finish. Controlled heating, superheat, and cooling rate wdl produce a uniform solidification fiont thus producing a uniform surface appearance and microstructure. In general a fine scan spacing was chosen with the power and scan speed being variables used to deliver different energy densities. The goal was to produce a short molten residence time with full through thickness melting.Prior process developmentg resulted in manufacture of test coupons (1 inch long, 0.5 inches wide and 0.060 inches thick) with acceptable microstructures for mechanical property characterization. The energy density used for producing these specimens was 3202 Jlcm2.

The energy density used for producing these specimens was 3202 Jlcm2. In order to develop acceptable processing parameters for the ABT geometry, a number of experiments were conducted investigating a wide array of energy densities. These tests proved that the ABT component having acceptable microstructure and surface finish could be produced in the 2400 ~ l c r n ~ range. The microstructure produced at 2400 J/cm2 is shown in Figure 3.to 3878 ~ l c m ~Figure 3. Microstructure of ABT produced at 2400 ~ / c m ~ , 50X (left) and 500X (right).Thermal Boundary EffectsExperimentation with different mold materials, energy densities, and preheats indicated that the solidification microstructure in a component can be controlled. This implies that the mechanical properties of a component can also be controlled or tailored through region specific manipulation of the microstructure.

This implies that the mechanical properties of a component can also be controlled or tailored through region specific manipulation of the microstructure. The variables listed above were manipulated to alter the cooling rate and thermal gradient and thus the solidification rate. The solidification dendrite arm spacing d can be described as a function of local solidification time t f or the average cooling rate6,, during solidification as1d = at.; = bk-" avgwhere a and b are constants, the exponent n is usually 0.3 to 0.5 for secondary spacingsand nearly 0.5 for primary spacings.The constants a and b are related because the local solidification time t f is a function ofthe average cooling rate i,, during solidification and the solidification temperature range AT, which is a physical property of the alloy. Thus,The average cooling rate in turn is the product of the temperature gradient at the solid-liquid interface G and the velocity of the solidification front R.

Thus,The average cooling rate in turn is the product of the temperature gradient at the solid-liquid interface G and the velocity of the solidification front R. Thus,It is evident that the temperature gradient and the solidification rate are easily manipulated through preheat, laser power density, scan speed, scan spacing and thermal boundary conditions at the mold-melt interface. Therefore the dendrite spacing as well as the solidification mode can be controlled through control of these parameters. For example, it has been shown9 that the solidification microstructure of SLS processed MAR-M-247 superalloy is a fbnction of laser energy density and thermal boundary conditions, ranging from fine-grain equiaxed to dendritic and directionally solidified. Mechanical properties such as fatigue can be improved through dendrite arrn spacing refinement.The solidification mode, i.e.

Mechanical properties such as fatigue can be improved through dendrite arrn spacing refinement.The solidification mode, i.e. plane front, cellular, dendritic or equiaxed as well as the types of defects that can form such as shrink and entrapped broken dendrite tips are known to depend on the cooling rate GR and the ratio G/R'~. Increasing GIR results in progressive transition from equiaxed to dendritic towards cellular onto planar front growth. Increasing the cooling rate GR results in finer structures. Thus GIR influences the type of structure whereas GR controls the scale of the structure.

Increasing the cooling rate GR results in finer structures. Thus GIR influences the type of structure whereas GR controls the scale of the structure. In laser surface remelting, the velocity of the solidification front is known to be related to the velocity of the moving heat source via the epitaxial growth angle"! Therefore, by appropriate control of process parameters in direct laser processing such as preheat temperature, laser power density, scan speed, scan spacing and thermal boundary conditions, it is indeed possible to locally control the type, scale and directionality of the solidification microstructure. An example of process control to obtain epitaxial growth is shown in Figure 4.Direct Laser Fabrication of a Gas Turbine Engine Component - Microstructure and Properties - Part IIProcessing AtmosphereBoth rough vacuum and the purged argon contained sufficient oxygen to oxidize the molten alloy thus causing balling, separation, and tearing due to surface tension effects.

The high vacuum atmosphere resulted in relatively uniform surface features, no cracking, tearing or separation when experimentation was conducted within an acceptable window of parameters. Volatilization of low vaporization temperature elements may be a concern depending upon preheat temperature and alloy composition. Volatilization can lead to problem with alloy depletion and or laser energy inconsistencies over long periods of time.Reactive metal coating thickness and qualityTwo vendors were evaluated for depositing the reactive metal coating on the abrasive ceramic grit. Significant differences were noted in how the product perfomed when subjected to the SLS process. In some instances the SLS preform would simply disintegrate while in other instances the preform was not sufficiently strong to withstand surface grinding. Internal contamination of the superalloy matrix was noted on the microstmcture of these specimens.

Internal contamination of the superalloy matrix was noted on the microstmcture of these specimens. In other cases the components perfomed flawlessly in all aspects. An analysis of the coating deposited on the abrasive grits explained the differences.Coated grit from two vendors were metallographically analyzed for coating thickness and uniformity. The photomicrographs shown in Figure 5 illustrate the difference in coating thickness and uniformity. The titanium coating thickness on Brand B is greater and much more uniform than of Brand A. Consequently, the Brand B material always performed better. However it was also noted that when the Brand A coating thickness was increased, the performance of Brand A still did not compare to that of Brand B. It was also noted that when only the Borazon 510 CBN (titanium coated cubic boron nitride) was used in the powder mixture the results were excellent.

It was also noted that when only the Borazon 510 CBN (titanium coated cubic boron nitride) was used in the powder mixture the results were excellent. Therefore, chemical analysis of the Ti coating on both Brand A and Brand B titanium coated alumina as well as the Borazon 5 10 CBN was performed. The contaminants C, 0, and N were of interest. Signifcant levels of these contaminants could dramatically diminish the performance of the Ti coating and consequently the overall integrity of the laser processed cermet.Two types of analyses were conducted in order to detemrine the contaminant content in the Ti coating. Auger analysis with intermittent ion milling from the surface of the particle provided data at different depths of the coatings. The data fiom this analysis is shown in Table 2 and Figure 5, Figure 6 and Figure 7. The tabular data at different depths indicates that the surface layers of all coatings analyzed were severely contaminated.

The tabular data at different depths indicates that the surface layers of all coatings analyzed were severely contaminated. A signifcant drop in co level was noted at 1000 A on all coatings, however the levels were still extremely high. A a large drop to relatively pure material was noted on the CBN coating while both the Brand A and Brand B dropped to near similar levels with Brand A having the highest overall cont level. At 5000 A only the CBN still possessed coating. The CBN's titanium coating at 5000 A was relatively pure. The data from the Auger Analysis indicated that the overall thickness and uniformity as well as purity at depths of approximately 1000 A is of high importance ine. Auger analysis could not discern nitrogen fiom titanium because of peak overlap.roprobe analysis was conducted on cross-sectioned samples.A t o d c % ConcentrationsCE3N Brand B - 5.2 wt %Brand A - 5.2 wt % Brand A - 2.4 wt. % # - Surface was charging too much to obtain data.

% # - Surface was charging too much to obtain data. This indicates that there was no free titanium or carbon * - Lack of data indicates that base particle had been exposed by ion milling. Note: Nitrogen cannot be discerned because nitrogen peak overlaps with titanium peak.Table 2. Auger chemical analysis of titanium coating.Titanium Content vs Coating Depth100.080 .0s 8 ,OeO 0 40.0 220.00.0020004000Depth in Angstroms= CBN I I -4- rand B - 5.2 wt. rand^ - 5.2-- - - - - .1%wt%i-BrandA-2.4wt %6000 I_j1Figure 6. Titanium coating purity vs. coating depth.-- --. - - - - --Oxygen Content vs Coating DepthIIi€I-- Brand B -+BrandA -I Figure 7. Oxygen content vs. titanium coating depth.0low20003m040W5000M)WDepth in AngstromsCarbon Content vs Coating DepthFigure 8 Carbon content vs. titanium coating depth.The results of rnicrobrobe analysis are shown in Table 3.

titanium coating depth.The results of rnicrobrobe analysis are shown in Table 3. The microprobe analysis also suggested that the CBN had the purest titanium coating. Brand A coating did have a much higher nitrogen content than Brand B with the overall contaminant content being similar. When a composite contaminant level is calculated for each of the abrasive grits using the C and 0 Auger data at 1000 Angstrom and the N levels from the microprobe analysis, a clearer picture of the performance is evident. The CBN composite contaminant content is approximately 28 atomic %, The Brand B Ti coated alumina is approximately 42%, While that of Brand A is approximately 63%. The chemical analysis data indicates that there is a significant difference in the coating purity of the three different grits. The results of the SLS trials combined with the chemical and metallurgical analysis indicated that the coating's uniformity and overall chemical purity are all equally important.

The results of the SLS trials combined with the chemical and metallurgical analysis indicated that the coating's uniformity and overall chemical purity are all equally important. The coating thickness is also important. However, the presence of significant residual coating on SLS processed specimens such as that shown earlier in Figure 3 indicate that the coating's thickness is not as important as uniformity or chemical purity.Microprobe Data for Cross - Sectioned Grit SamplesTable 3. Microprobe data for cross-sectioned grit samples.Effect of DopantsDue to some of the processing difficulties encountered because of the coating uniformity and purity, an experiment involving doping was performed. Doping consists of adding a material to the cermet mixture that would assist in scrubbing and breaking down the highly contaminated surface of the Ti coating on the grit particle and also assist in cleaning the remainder of the melt.

Scmbbhg of the Ti coating is necessary in order to clean the refractory type surface and activate the more pure underlying Ti in order to improve the wetting and bonding with the superalloy matrix. Different methods were evaluated for adding various levels of free Ti to the melt. Other elements that have been evaluated as dopants are F and Y. Both methods yielded promising results. In addition to Ti doping, multiple coatings were also evaluated. The coating in this caseStress-rupture and rub-testing of SLS processed ABT cermet material have been reported The stress rupture samples were divided into two distinct groups: ABT specimens with uniform surface appearance and ABT specimens with some "acceptable" levels of pitting andporosity. The surface condition did not produce any difference in stress mpture properties.

The surface condition did not produce any difference in stress mpture properties. An additional variable involved el ating the 2 1 50°F/32 hour standard diffusion cycle from some of ion of the braze diffusion heat treat cycle significantly improves the the specimens. -4 base material. The stress-rupture data in Larson Miller form is shown properties of th in Figure 10. The testing was conducted in the temperature range of 1900°F to 2200°F with some of the samples tested in excess of 500 hrs. The samples brazed without the diasion heat treat cycle exhibited slightly reduced life. All samples failed in the braze joint and did not traverse the SLS material. This indicates that the plot of Figure 1 0 is actually a representation of the braze bond strength. Standard production material exhibits failur traverse through the center of the cemet because of the braze component in the standard e.

Stress-Rupture of CMSX-4 bar with ABT brazed in-gage.The primary goal of future work is to develop a laser process for direct tipping. Direct tipping is perfo SLS on top of the turbine blade and actually bonding the cemet layer to the blade at the same time that melting and densification is occuning. A laser based process that fabricates abrasive tips directly onto turbine blades will yield three important benefits. The first and most important potential benefit is that the base material of the blade will have substantially better mechanical properties since it will not be exposed to either the braze or diffusion thermal cycles. Both the braze and diffusion cycles involve high temperatures and long times which degrade the microstructure and mechanical properties of the base material. cycles will allow higher design mechanical propertiestemperatures, potentially resulting in designs that improve eficienc y.

cycles will allow higher design mechanical propertiestemperatures, potentially resulting in designs that improve eficienc y. Second, the need fortooling and the intensive labor associated with preparing and brazing tips to blades will be eliminated. Thrd, mechanical properties of the tip-blade bond interface will be improved as a result of epitaxial solidification of the tip material off the blade base material.Additional goals of the fiture include incorporating process understanding and control such that region specific or tailored microstmctures can be produced. This ability could significantly improve the performance of a given region or component overall thus increasing expected life. Such achievements may also allow the repair of single crystal turbine hardware to original conditions or the production of super plastically formable components of a given alloy which were not available previously.

The total control over processing environment should favor technological breakthroughs in the manufacture of net shape components in extremely difficult to process high performance materials such as refractories, in-situ composites, and intermetallics.CONCLUSIONSA successfil direct SLS process was developed and approximately 100 tips produced. Theprocess developed achieved all of the objectives and goals originally set forth. The goals are listed below .1. Raw material input is proportional to quantity of tips needed.2. Fall off materials are reusable.3. Majority of processing steps of standard production process eliminated.4. Process produces the hlly dense component in a single pass by achieving hll through thickness melting.5. Mechanical properties were improved by eliminating the cobalt braze material from the cermet composition .6.

Mechanical properties were improved by eliminating the cobalt braze material from the cermet composition .6. Elimination of Processing steps, scrap, fall off, and input stock required produced cost savings in excess of that originally calculated.7. Microstructure, properties and performance are superior to that of the standard production material.In addition to achieving all of the original program goals and outlining parametric windows, important information was noted concerning the effects of preheat temperature, processing atmosphere, thermal boundary effects, energy density, dopants, and reactive metal coating quality on ceramic grit in cermets. Some of these effects are summarized as follows.Preheating the powder bed has been shown to be beneficial by outgassing the powder thus preventing firther contamination during processing.

Some of these effects are summarized as follows.Preheating the powder bed has been shown to be beneficial by outgassing the powder thus preventing firther contamination during processing. Preheat was also noted to improve the uniformity of flow and solidification of the molten material, thus enhancing the surface finishuniformity of the final component. It is also known that preheat to appropriate temperatures also reduces hot tearing, hot cracking, and balling of the molten material.Energy density and the associated parameters of laser power, scan spacing, and scan speed were crucial in minimizing hot tearing, balling and cracking. Without the proper selection of energy density parameters significant porosity, incomplete through thickness melting, and very poor surface fimsh may result.The SLS apparatus developed during this program allows total process control.

The extent of process control capability allows the manipulation of thermal boundary effects and solidification rates. This in turn may be used in the future to tailor region specific microstructures. Experiments conducted during this program illustrate that it is possible to dramatically manipulate the microstructure of the material.The experimentation with different types of abrasive grit, Ti coatings and contaminant analysis revealed the importance of uniformity and purity of the reactive metal coating in the processability, performance and properties of cermets produced by direct SLS. High quality coatings can result in the difference between failure and success.Dopants were evaluated in an attempt to overcome the detrimental affects of the poor quality reactive metal coatings. The usefulness of dopants to improve melt cleanliness and activate the reactive metal coating on the abrasive grit to enhance bonding was demonstrated.

The usefulness of dopants to improve melt cleanliness and activate the reactive metal coating on the abrasive grit to enhance bonding was demonstrated. However, caution should be exercised with dopants due to secondary affects resulting from alloy chemistry modification. The coated grits used for production were of sufficient quality that a dopant was not necessary to provide a robust production process. However, dopants may be very beneficial if one were to produce a very large part or multiple layer part.ACKNOWLEDGEMENTSPartial funding for this work was provided by the United States Air Force, Contract No. F336 15-94-C-2424. The remainder of resources were provided by Allison Engine Company and the University of Texas at Austin.1 Marcus, H.L and D. L. Bourell, Solid Freeform Fabrication Finds New Applications, Advanced Materials & Processes, Volume 144, No.3, September 1993.

RECT LASER FABRICATION OFPERFORMANCECOMPONENTSVIA SLS/HIPMartin Wohlert, Suman Das, Joseph J. Beaman, andDavidBourellLaboratory for Freeform Fabrication The University of Texas at Austin ETC 5.160 C2200 Austin, TX 78712-1063AbstractRecent research in the area of direct freeform fabrication of components via selective laser sintering/hot isostatic pressing (SLS/HIP) has focused on the processing of Alloy 625. Alloy 625 is a nickel-based superalloy which provides high temperature strength and corrosion resistance. Alloy 625 test specimens were successfully SLS processed with an integral gas impenneable skin or "can". These samples were subsequently HIPed to high density (>99.5%). Characterization of the test specimens indicated that microstructures similar to conventionally processed P/M materials are achieved in the HIP consolidated "core" region of the parts, while structures similar to those found in cast materials are present in the SLS processed "can" regions.

Mechanical analysis of Alloy 625 SLS/HIP parts and production of complex structures will commence shortly.Introductionshape manufacturing technique under development for the production of high performance metal components 1 • SLS/HIP combines the freeform shaping capability of selective laser sintering (SLS) with the full densification capability of hot isostatic pressing (HIP) to produce net shape, high value metal components with significantly reduced costs and lead times in comparison to traditional HIP processing. Components being evaluated for this process include aircraft turbine engine hardware, naval, and submarine components.Selective laser sintering/hotisostatic pressing (SLS/HIP)is a hybrid netA schematic of the SLS/HIP process is shown in Figure 1. As an object is built layer by layer, the laser fuses the metal powder only at the part boundaries to fonn an integral, gas impermeable skin or "can" to a density exceeding 92% theoretical density.

As an object is built layer by layer, the laser fuses the metal powder only at the part boundaries to fonn an integral, gas impermeable skin or "can" to a density exceeding 92% theoretical density. This is the fractional density at which the porosity typically changes from interconnected to closed2 • The powder in the interior of each layer cross-section can be left unprocessed or optionally laser sintered to an intermediate density. Thus, the component is shaped, canned, evacuated and sealed in-situ. The integrally canned, net shape component produced by SLS is directly post-processed by containerless HIP to full density. A final machining step may be applied if necessary.281Laser beam seans high density skin for cnrrent cross-sectionLaser beam scans interior of current cross-sectionIn-situ canningSLS/HIPDirect SLStFinal MachiningHot Isostatic PressinFigure 1: The SLS/HIP process.SLSIHIP has several advantages over conventional HIP methods.

Since an integral skin or "can" is formed of the same material as the component, a secondary canning step is not necessary. Adverse container-powder interactions are eliminated and post-HIP container removal is not required. Tooling and pre-processing steps associated with container fabrication and filling are also eliminated. Therefore, SLSIHIP enables production of complex shapes at reduced cost and in shorter lead-times.is being undertaken for two materials and two demonstration components. Based on a surve/ of several naval installations, the materials selected for technology development and demonstration are Alloy 625 and Ti-6AI-4V. Prior work on SLSIHIP development for Ti-6AI-4V has been reported elsewhere3As part of an ongoing research program, SLSIHIP process developmentThis paper focuses on process development for Alloy 625.ExperimentalAlloy 625 is produced by several manufacturers, under a variety of trade names.

The fundamental properties and chemical composition are available in test standards such as ASTM B443, B444, and B446. The composition and metallurgy of Alloy 625 is similar to the more commonly used Alloy 7184 • Its notable properties include high tensile and fatigue strength, as well as excellent corrosion and oxidation resistance. The corrosion resistance is largely provided by the high chromium content of the alloy, while creep strength is provided by formation of a variety of carbides and precipitates (primarily "("). However, the precise quantity and282distribution of microconstituents can vary greatly depending on the exact composition and thermal history of the material. The supplier selected to provide a powder form of the material was Anval Inc. (301 Route 17 N., Suite 800-T, Rutherford, NJ 07070). The chemical composition and particle size distribution for the supplied powder is listed below.. 1 Table 1.

(301 Route 17 N., Suite 800-T, Rutherford, NJ 07070). The chemical composition and particle size distribution for the supplied powder is listed below.. 1 Table 1. C emica ComposItionhNiCr MoFeNbTiAlSiN Mn CoCS62.221.38.224.003.360.270.210.16.051.04.04.028.008Table 2. Particle Distribution Microtrac)Micron4431221611%<78.840.69.42.61.5Analysis of the provided powder by SEM and optical microscopy indicates a generally spherical morphology with low internal porosity, as shown in Figure 2.Figure 2. SEM micrograph of spherical Anval 625 powder (left). Optical Micrograph of polished powder showing low internal porosity (right).Initial experiments with Alloy 625 were designed to demonstrate the feasibility of constructing integrally canned shapes by SLS suitable for HIP post-processing. To this end, a number of cylindrical specimens were fabricated by SLS.

To this end, a number of cylindrical specimens were fabricated by SLS. Processing was performed in vacuum using a 250 watt continuous wave Nd:Yag laser. Early test samples showed a high degree of porosity and cracking after processing. The presence of cracks or connected porosity cannot be tolerated in the walls of the part since a gas tight specimen is critical for successful HIPing. Further analysis of the problem indicated that contamination of the powder was a likely culprit. The problem of absorbed gases and other contaminates on powder is well known in traditional powder metallurg/. The removal of contaminants can be accomplished by "degassing" the powder in vacuum. Typically, absorbed water vapor is found in the greatest quantities. The283P.007application of heat to the powder can greatly accelerate the rate of contaminant removal. Measurement of contaminant concentration versus temperature by residual gas analysis (RGA) was performed.

Measurement of contaminant concentration versus temperature by residual gas analysis (RGA) was performed. Scans for a typical metal powder are shown in Figure 3. A temperature increase as low as 150°C is sufficient to greatly accelerate contaminant removal. Removal of absorbed gases produced a dramatic improvement in alloy flow and densification, as shown in Figure 4...;..>ib 100 ] 90 ~ 80 0.. 70 t ~ 60 ~ ~ 50 S 40 ~ 30 ~ 20 ~ 10 ~(J)~50100 150200 250400 Temperature (OC)300350450500550600650700750Figure 3. Residual Gas Analysis (RGA) data showing water desorption from metal powder with . • 6 IncreaSIng temperature .Figure 4. As-SLS processed specimens, contaminated powder (left), after 12 hour vacuum bakeout at 450°C (right).Following optimization of the processing parameters, production of cylindrical HIP specimens commenced.

A typical specimen, shown in Figure 5, has an outer diameter of 0.5 inches with a wall thickness of 0.125 inches and an unprocessed powder core. A portion of a transverse cross-section of one such cylinder is shown in the photomicrograph of Figure 6. The figure shows a section of the powder core, the right side-wall and the base of the cylinder. The284fractional density varies from 60% at the axis to1110rethan 98.5% at the side-wall. Shown on the left in Figure 7 is iaphotomicrograph of the interface betweenthecylinderside-",all and the interior. powder core while the photomicrograph on the right shows.theinterfacebet'Ye.en the fully dense side-wall and the porous outer layer formed by sintering due to local heatttansfer.Figure 5. A typical as-SLS processed Alloy 625 cylinder.~' . ~..........,!,o ...."'.!)';.. >,..."',"'.Figure 6.

A typical as-SLS processed Alloy 625 cylinder.~' . ~..........,!,o ...."'.!)';.. >,..."',"'.Figure 6. Transverse cross-section segment of a typical SLS processed Alloy 625 cylinder at the interface betweenbpttom-ca.p, side-wall and core.285MicrostructureSample parts were sectioned, mounted, and polished for metallographic inspection. Grain structure was revealed by a mixed acids etchant7 (15 ml HCI, 5 ml acetic acid, and 5 ml HN03). The photomicrograph on the left in Figure 8 shows the grain structure of as-SLS processed fully dense skin in a cross-section transverse to the powder layer orientation. The image height is the equi~alent of approximately 7 powder layers, each of thickness 125JlID. The microstructure conSIsts of elongated columnar grains oriented vertically in the build direction.

The image height is the equi~alent of approximately 7 powder layers, each of thickness 125JlID. The microstructure conSIsts of elongated columnar grains oriented vertically in the build direction. The grain orientation indicates that the remelted surface of the previously solidified layer, which provides significantly better heat conduction than· the loose powder surrounding the part along the sides. No distinct boundaries between individual powder layers are visible. The photomicrograph taken parallel to the powder layers exhibits the expected equiaxed grain structure, as shown on the right in Figure 8.the solidification front originates atFigure 8. Etched microstructure of cylinder side-wall, transverse section (left) and longitudinal section (right).cycle: ramp from ambient temperature to 1150° C at 15° Clmin, pressurize to 138 MPa (20,000psi) over 30 min, hold at 1150° C and 138 MPa for 3 hours, ramp down to ambient at 30° C/min.

Image Figure 9 shows photomicrographs of the powder core taken prior to and after HIP. analysis of the unprocessed core indicates a density of 60% prior to HIP. Actual density may be slightly higher, in the range of 65%, due to errors caused by particle pullout during sectioning and polishing of the sample. Following HIP, a core density of 99.5% was achieved. Etching reveals grain boundaries formed largely along prior particle boundaries. The generally spherical geometry of the particles is preserved in the grain structure shown in Figure 9. While this is the expected microstructure for HIP processed parts, the effect of the prior particle boundaries (PPBs) on the mechanical behavior of the parts has not yet been established. The presence of contaminants or undesirable precipitates at PPBs can substantially reduce ductility and tensile strength.

The presence of contaminants or undesirable precipitates at PPBs can substantially reduce ductility and tensile strength. Because there is no additional cold working to disrupt the boundaries followin~ HIP, the presence of a continuous impurity film can be quite detrimental. Previous researchers have identified the presence of an AI-rich oxide film as limiting mechanical and corrosion properties in improperly processed material. While great care has been taken during SLS processing to minimize the possibility of contamination, tensile test specimens will be produced in the nearSLS processed Alloy 625 cylinders were post-processed using the following286future to verify the mechanical properties of the SLSIHIP processed material.Figure 9. Unprocessed powder core, approximately 60% dense (left), etched microstructure following99.5% dense (right).A second type of cylinder was also processed by SLS. This cylinder, fabricated with an identical 0.5 in.

This cylinder, fabricated with an identical 0.5 in. diameter, was processed to near full density across the entire diameter. This type of processing is necessary to fabricate "end-caps" for an integrally canned object made by SLS the density in this type of suitable for specimen exceeded 98.5% of theoretical density. The etched microstructure, shown on the left in Figure 10 reveals equiaxed grains. This microstructure is quite comparable to that obtained in conventionally cast or annealed wrought materials9 • The SLS processed specimen was post processed by a HIP cycle consisting of 3 hours at 1240°C and 172 MPa (25000 psi). The microstructure in the HIP post-processed specimen (with density exceeding 99.5% of theoretical) is shown on the right in Figure 10.post-processing. Metallography revealed thatFigure 10.

The microstructure in the HIP post-processed specimen (with density exceeding 99.5% of theoretical) is shown on the right in Figure 10.post-processing. Metallography revealed thatFigure 10. Etched microstructure of 98.50/0 dense SLS processed (left) and 99.50/0 dense SLS/HIP processed (right) Alloy 625.287ConclusionsFeasibility of fabricating integrally canned complex shapes by SLS for containerless HIP post-processing was demonstrated in Alloy 625. Simple cylindrical shapes were fabricated by SLS and consolidated to full density by HIP. Preliminary evaluation of microstructure reveals that material processed by the SLS/HIP technique is equivalent to conventionally processed P/M material in the part core. The small volume of material processed directly by SLS that forms the can structure is similar to conventional cast structures.

The small volume of material processed directly by SLS that forms the can structure is similar to conventional cast structures. Future work will focus on building a part with complex geometry as a technology demonstration for Alloy further characterization of structure/property relationships as well as optimization of processing parameters for build speed, skin thickness and HIP cycle time.AcknowledgementsThe authors wish to acknowledge research funding provided by DARPA and ONR undercontract NOOOI4-95-C-0139 titled "Low Cost Metal Processing Using SLS/HIP".ReferencesI R. Knight et aI., "Metal Processing Using Selective Laser Sintering and Hot Isostatic Pressing", Proc. Solid Freeform Fabrication Symposium 1994, (Austin, TX: U. of Texas, 1994). 2 R. M. German, Powder Metallurgy Science, Chapter 7, pp. 257, MPIF, 1994. 3 S. Das et aI, "Producing Metal Parts by Selective Laser Sintering/Hot Isostatic Pressing", JOM, December 1998, TMS. 4 S.

RECT LASER SINTERING OF BOROSILICATE GLASSF. Klocke, A. McClung and C. Ader Fraunhofer Institute for Production Technology IPT, Aachen, GermanyReviewed, accepted August 4, 2004AbstractDespite the advantages that selective laser sintering (SLS) offers in terms of material availability, many materials have yet to be explored for feasibility and even fewer are available on a commercial basis. This paper presents initial investigations for one such material, borosilicate glass, which could be of particular interest to filter manufacturers because it presents an attractive alternative to the conventional, time-consuming way of producing filters of various porosity classes. Process results are presented including a determination of the optimal parameter window and the effect of processing parameters on the density and surface quality.

Process results are presented including a determination of the optimal parameter window and the effect of processing parameters on the density and surface quality. The effects of thermal post-processing and the inclusion of an additive are also discussed.IntroductionIn order for new applications to emerge in solid freeform fabrication industry, materialsneed to be developed that match those used in conventional machining processes. However, the applicability of the material for the rapid process must be ensured. SLS in particular needs to be exploited for new material investigations, due to the fact that any material that can sinter or has a decaying viscosity curve can theoretically be employed. The borosilicate glass powder explored herein is one such material, which has potential interest to manufacturers of porous borosilicate glass products, such as filters and sensors.Borosilicate glass can be classified as a standard glass.

Thereby, by definition, it has theproperty that it exists in a vitreous state in which there is no long range atomic order. The glass powder employed herein is borosilicate glass type 3.3 with an average particle size of 30 µm and a broad distribution range. The composition is largely SiO2 mixed with B2O3, Fig. 1. The presence of B2O3 reduces the coefficient of thermal expansion, to about one-third that of standard soda-lime glass, and makes the glass more resistant to thermal shock [1]. As a glass, it has no exact melting point but rather is defined by its softening point, 820 °C, and by the glass transition temperature, Tg = 552 °C. When heated above Tg, the glass properties of borosilicate transform into melt properties.

Therefore, first the three main parameters were explored in single-layer trials: PL, vs and hs in order to determine the sinter depth and behavior. The single layers are 20 x 20 mm2 and were exposed in a 2 mm deep powder bed. A target value of ds = 0.2 mm was chosen. Clearly ds needs to be larger than the largest particle size and sufficiently large to provide sufficient adhesion to the previous layer. However it should also be kept relatively small in order to inhibit the thermal stresses resulting from the re-exposure of previous layers.The initial results for borosilicate glass show the same trends as is known from ceramic SLS processes [3]. Figure 3 shows a summary of the single layer investigations in terms for ds versus PL, vs, and hs. An increase in PL led to an increase in ds, as shown in the top right graph for increasing PL with hs = 0.12 mm and vs = 500, 1000 and 2000 mm/s.

An increase in PL led to an increase in ds, as shown in the top right graph for increasing PL with hs = 0.12 mm and vs = 500, 1000 and 2000 mm/s. As anticipated, ds also increases as vs decreases, as shown in the bottom left graph for PL = 14 W and hs = 0.03, 0.12, 0.21 and 0.3 mm. As hs is increased, ds decreases within the range shown on the bottom right hand graph for PL = 14 W and vs = 100, 200 and 300 mm/s. From these results, parameter sets can be combined. For instance, for dn = 100 µm, PL = 14 W, hs = 0.12 mm, and ds = 0.23 mm, the optimal scan speed is approximately vs = 400 mm/s.Cubes, a multilayer test geometry, were built in order to provide a more accurate analysis in terms of visual analysis, surface roughness and density. These cubes are 20 x 20 x 10 mm3. This is a sufficient height for a visual analysis of side surface roughness and cracks. With these cubes, other parameters were then explored, such as dn and the exposure strategy.

This is a sufficient height for a visual analysis of side surface roughness and cracks. With these cubes, other parameters were then explored, such as dn and the exposure strategy. The main paramters were varied from PL = 5 – 94 W, vs = 100 – 1300 mm/s and hs = 0.03 – 0.30 mm. Parameter sets were matched from the single layer exposures for ds ≈ 0.2 mm. dn = 100 µm and a layerwise alternating unidirectional exposure strategy was employed.

Parameter sets were matched from the single layer exposures for ds ≈ 0.2 mm. dn = 100 µm and a layerwise alternating unidirectional exposure strategy was employed. The base plate used was of the same material composition to obtain better adhesion to the sintered part.215Material Borosilicate glass 3.30.70.6ds as f(PL)0.5Machine Parameters PL = 5 - 94 W vs = 50 - 3000 mm/s hs = 0.03 - 0.3 mm]m m[sd0.40.30.20.1Geometry 20 x 20 mm2, single layer002040P L [W]6080vs:500 mm/s1000 mm/s2000 mm/s1ds as f(vs)0.8ds as f(hs)0.80.70.6]0.6m m[]m m[0.50.40.4sdsd0.30.20.20.100010020030040050060070080000.050.10.150.20.25vs [mm/s]hs [mm]hs:0.03 mm0.12 mm0.21 mm0.3 mmvs:100 mm/s200 mm/s300 mm/sFig. 3: Single layer experimental results – dependence of parameters on dsCrack propagation was common for all parameter combinations.

3: Single layer experimental results – dependence of parameters on dsCrack propagation was common for all parameter combinations. These crack profilesare visible over a number of layers, unlike with delamination. One theory for this is the thermal shock characteristics of borosilicate glass. Borosilicate glass, while more resilient than many glasses, is still known to have problems with accommodating rapid temperature changes of greater than 5 °C/min [2]. The laser beam clearly introduces more heat in a shorter period of time, which may lead to thermal stresses within the sintered part and be a primary factor in the resultant crack profiles. Over 47 W, these cracks led to failure in the parts. Therefore the experiments were found to be only successful at low PL and thereby low vs. Figure 4 shows the results for PL = 14, 32 and 47 W consecutively. It is visibly clear that at 14 W, the appearance of cracks is greatly reduced and the stability of the parts thereby improved.

Figure 4 shows the results for PL = 14, 32 and 47 W consecutively. It is visibly clear that at 14 W, the appearance of cracks is greatly reduced and the stability of the parts thereby improved. Hence PL = 14 W was taken as an upper limit for further investigations.Fig. 4: Test geometry experimental results for increasing PLNext hs was varied over the range from 0.03 mm – 0.30 mm with a constant PL = 9 Wand ds ≈ 0.2 mm. While parts could be built over the entire range, the side surface appearance at hs ≤ 0.12 mm was extremely poor. In Figure 5 the differences are clearly visible: the left part was produced with hs = 0.03, compared to the right part with hs = 0.18 mm, where the side surface is greatly improved. Furthermore, an upper limit on hs was determined to be at2161000.3approximately 0.30 mm, because the top surface roughness was highly degraded above this value.Fig.

Furthermore, an upper limit on hs was determined to be at2161000.3approximately 0.30 mm, because the top surface roughness was highly degraded above this value.Fig. 5: Test geometry results for varied hsTwo main analysis parameters, the density and surface roughness Rz, were thenconsidered for a range of parameters as shown in Figure 6. Again the same basic mechanisms were found to prevail. An increase in vs leads to a decrease in the density of the specimen, shown in the top left graph. However, Rz also increases, as shown in the top right graph. Therefore choosing between higher density or lower Rz will still be a tradeoff when choosing an optimal vs. The results for hs, bottom left, show on increase in density at lower hs. However the density seems to taper off at hs = 0.18 mm, and furthermore Rz shown on the bottom right, increases drastically above 0.18 mm, leading to the conclusion that there is no advantage to having hs ≥ 0.18 mm.

Furthermore, prior results indicated that the side surface was greatly degraded below hs = 0.12 mm, resulting in a narrow window of 0.12 ≤ hs ≤ 0.18 mm.60,00Density as f(vs)200Rz as f(vs)]%. l o V[y t i s n e D55,0050,00]m u [zR18016045,00140100150200250300100150200250vs [mm/s]vs [mm/s]54,50Dens ity as f(hs)220R z as f(hs)]%. l o V52,50]200[y t i50,50m u [180s n e D48,50zR16046,500,060,12hs [mm]0,180,241400,060,12hs [m m ]0,18Fig. 6: Density and surface roughness Rz for increasing vs (top) and hs (bottom)A secondary consideration was the dimensional accuracy of the parts. All of the parameter combinations which led to a successfully sintered part were slightly larger than the anticipated size (approx. 0.30 – 0.60 mm along each length). The most likely cause is the heat transferred by the laser beam.

0.30 – 0.60 mm along each length). The most likely cause is the heat transferred by the laser beam. The sintered region differs from the region traversed by the center of the laser beam, resulting in slightly larger parts. Part growth is also caused by heat2173000,24transfer from the exposed part to the surrounding loose powder. These effects can however be compensated for with a beam compensation factor.Investigations into the implementation of other exposure strategies, such as the stripesand chess exposure strategies, were also held. Both strategies benefit from shorter scan vector lengths, but often with a sacrifice in surface quality and part strength. For the employed geometry, a jump in density was seen with a 5 mm wide stripe or chess width. For the chess strategy, the resulting density, 49.5 %, was nearly a percent larger than with a unidirectional exposure strategy.

For the chess strategy, the resulting density, 49.5 %, was nearly a percent larger than with a unidirectional exposure strategy. However part strength was also greatly reduced as the parts were prone to break along the fill lines, or overlap, between chess squares. Furthermore the optimal stripe or chess size is largely dependent on the given part geometry.Of all the parameter combinations that led to a stable laser sintering process, themaximum density achieved was 48.6 % of theoretical density (PL = 9 W, vs = 225 mm/s, hs = 0.15 mm) with Rz = 152 µm.To examine the microstructure and better understand the density achieved, the sintered parts were then subject to a post-heat treatment in an oven for 6 hours at temperatures of 500 °C to 900 °C. As expected, there was no significant change in properties at 500 and 550 °C.

As expected, there was no significant change in properties at 500 and 550 °C. At 600 °C, however, there is a small increase in the density (slightly less than 1 %), signaling that only at this point is the rearrangement of boron oxide, with a softening point of 450 °C, complete. At 700 °C, the part begins to shrink (3.1 %), but there is a significant increase in density from 49.7 % to 54.4 %. This trend continues at 750 °C where the part starts to deform and has shrunk an additional 5 %. At 800 °C, there is significant shrinkage (17 %) and warpage, but the part reaches nearly full-density and at 900 °C the part is no longer identifiable. This is consistent with the softening point of borosilicate glass at 820 °C. The SEM figures show correlating results. At 600 °C, there is much unsintered powder, however at 700 °C, more sintered structures are formed, and from 750 °C – 800 °C the pores close off to achieve near full density.

At 600 °C, there is much unsintered powder, however at 700 °C, more sintered structures are formed, and from 750 °C – 800 °C the pores close off to achieve near full density. These results are summarized in Figure 7.600 °C700 °C750 °C800 °CTest GeometryMicrostructureρ (%) Shrinkage (%)49.7 054.4 3.163.3 ≈ 8.096.4 ≈ 17.0Fig. 7: Post-processing of laser sintered borosilicate glassTo determine if this density could be increased with the use of an additive, 0.05 %weight carbon black powder was mixed with the borosilicate glass powder. During the manufacturing of carbon black, aggregates are formed of primary particles which are only a few tens of nm in size. Due to the small size and the small distances between the aggregates, van der Waals forces are present, and as a result the aggregates cluster to form agglomerates.

Due to the small size and the small distances between the aggregates, van der Waals forces are present, and as a result the aggregates cluster to form agglomerates. When mixed with borosilicate, the agglomerates break up to form a network of fine particles218with thousands of cubic meters of carbon black on the surface throughout the matrix. Thereby carbon black does not sinter but provides higher energy absorption sites allowing the borosilicate glass to more tightly absorb.The addition of this powder increased the sinter depth of the mix, enabling theborosilicate glass to be sintered at slightly higher temperatures. There was however no significant increase in the viable parameter window. Comparing the results of a stable parameter set for both with and without carbon black, the carbon black led to a comparable surface roughness and a small increase in density of 0.8 %, Figure 8.

Comparing the results of a stable parameter set for both with and without carbon black, the carbon black led to a comparable surface roughness and a small increase in density of 0.8 %, Figure 8. Furthermore the use of carbon black necessitates the use of a post-processing step for removal.BSCG + 0.05 % Carbon BlackOnly BSCGTest GeometryMicrostructureρ (%)49.548.6Fig. 8: Effect of carbon black additive on borosilicate glass SLS partConclusionBorosilicate parts could successfully be produced using SLS within the parameter window: PL = 5 – 9 W, vs = 300 – 550 mm/s and hs = 0.12 – 0.18 mm. One stable parameter combination was found to be: PL = 9 W, vs = 225 mm/s, hs = 0.15 mm and dn = 0.10 mm. This resulted in a density of 48.6 % of the theoretical density, an Rz of 152 µm and a dimensional accuracy of 98 %. The major drawback was the appearance of crack profiles, which is assumed to be due to the thermal shock characteristics of borosilicate glass.

The major drawback was the appearance of crack profiles, which is assumed to be due to the thermal shock characteristics of borosilicate glass. One likely way to alleviate this problem is with the implementation of a heated powder bed in order to reduce the thermal gradients. A thermal post-processing at 700 °C results in a slightly denser part, 54.4 %, with a reduced appearance of cracks due to mechanism of viscous flow.References[1] Doremus, R.H. Glass Science. Wiley-Interscience (1994)[2] Robu Glasfilter-Geräte. Sinterfilter Vitrapor: Technical Data.http://www.robuglas.com/[3] Klocke, F., Wirtz, H. Selective Laser Sintering of Zirconium Silicate. Proceedings of the9th Solid Freeform Fabrication Symposium, Austin, pp 605-612 (1998)219.

RECT LASER SINTERING OF CERAMICSF. Klocke, C. Ader Fraunhofer Institute for Production Technology IPT, Aachen, GermanyAbstractFor more than one decade layer manufacturing technologies assist the development of new products. Due to a layer-wise build-up of a three-dimensional geometry, nearly every complex design is producible in a short period of time. Selective Laser Sintering is a powder- based technique to produce plastic prototypes (Rapid Prototyping) or metal mould inserts (Rapid Tooling). The laser sintering of ceramic powder is not yet commercialized but applications could be both Rapid Prototyping and Rapid Tooling. The former involves the laser sintering of investment casting shells and cores to cast metal prototypes and the latter the laser sintering of ceramic master patterns for metal spray forming of steel mould inserts. The advantage compared to actual processes are a faster availability of the final product.

The advantage compared to actual processes are a faster availability of the final product. To facilitate these applications, special ceramic powders as well as new process parameter combinations were investigated. This paper will present achieved results within the above- described applications.IntroductionSteadily decreasing product life cycles make it ever more important to reduce the development time for new products. Time-to-market instead of development costs has become the key factor for a products success. New methods from the organizational (Simultaneous Engineering) and the operative point of view have been developed [1].At the end of a product development, technical prototypes which closely resemble the final product are demanded. Viable solutions for plastic components have been around for years, but the prototyping of metal components has often thought to be too time-consuming, and has been unable to meet comparable specifications.

The fast production of metal prototypes still remains a central concern in the area of layer manufacturing technologies [2].Within the sphere of direct manufacture of metal prototypes, a few technologies suchas 3D Printing or laser sintering are already commercialized. Since laser sintered or 3D printed metal prototypes contain low melting alloys, the mechanical properties are limited.The indirect manufacture of metal prototypes provides two different approaches either the direct Rapid Tooling where mould inserts instead of prototypes are layer-manufactured or the combination of Rapid Prototyping techniques and casting processes (indirect Rapid Tooling). The direct Rapid Tooling is utilized if higher quantities of prototypes are necessary. To date, this application achieves good results for plastic injection molding.

The direct Rapid Tooling is utilized if higher quantities of prototypes are necessary. To date, this application achieves good results for plastic injection molding. Examinations and case studies of Magnesium and Aluminum die casting or thixo casting are carried out at several research institutes [2, 3, 4].447If smaller amounts of metallic prototypes are required the combination of rapidtechniques and casting processes such as investment casting is used. Virtually, all metals may be cast and thin walls or small details are readily reproduced because the metal is poured into a closed hot shell. With actual Rapid Prototyping techniques the manufacturing time for the necessary master pattern is minimized but the time-consuming slurrying, sanding and drying of the samples is not eliminated.Ceramic Laser SinteringThe laser sintering of ceramic materials is comparable to plastic or metal laser sinteringwhereas ceramic laser sintering is a transient liquid-phase sintering.

With regard to the process, the combination of layer thickness, laser power, scan velocity, hatch spacing and scanning vector length has to be adjusted to the ceramic powder, Fig. 1.Fig. 1: Process parameters of ceramic laser sinteringThe investigations on ceramic laser sintering at Fraunhofer IPT started on a laser sintering system with up to 100 W of laser power (CO2). From initial tests with different ceramic powders such as aluminum oxide, aluminum silicate or zirconium silicate, the zirconium silicate powder (ZrSiO4) was found the most suitable to process [5]. Today, a 200 W laser sintering system is used.The basic research with zirconium silicate was carried out with ground powder. Several grain size distributions like 125 mesh, DIN 70 and 325 mesh and mixtures of these powders were tested.

Several grain size distributions like 125 mesh, DIN 70 and 325 mesh and mixtures of these powders were tested. The result was that a high portion of small grains not only improves the surface quality but also improves the sintering itself since smaller particles need less energy to melt than larger particles. In contrast, an increased amount of small grains increases agglomeration so that an automatic recoating with a uniform surface was not possible. Dedusting and fractionating of the powder solved these problems. Today, the use of granular powder is preferred: due to the spherical surfaces a recoating with small granules is possible.To determine the influence of process parameters on laser sintering each parameter has to be varied on its own. An increase in laser power PL leads to an increase in the density of the specimen, since more energy is delivered into the powder and larger melt pools fill up the porous structures of previous layers.

However, surface roughness also increases because the448molten particles tend to form larger spherical structures due to their effort to reduce the free enthalpy by optimizing the ratio between the area of free surfaces and the related volume.Increasing the scan speed vS has the opposite effect: due to higher velocity of the laser focus, less energy is delivered into the powder bed and less material is sintered. The result is that the density decreases but the surface finish improves.A variation of hatch spacing hS leads to different effects: if hatch spacings are too small,the molten powder particles form larger beads of high density leading to an overall high density of the part, yet the surface roughness also increases.

For a hatch spacing larger than the focal diameter, the molten beads of material from line to line are not connected to each other and the gaps are filled with unsintered powder, resulting in a rough surface and low density respectively low strength.A reduction of layer thickness dn allows a faster scan velocity because the necessary sintering depth to ensure a fusion to the previous layer is reduced. Additionally, a smaller layer thickness increases the dimensional accuracy of the laser sintered part since the step effect is minimized. The thickness is limited by the grain size of the powder. Actually, the smallest layer is 50 µ m.Of all the parameter combinations that lead to a stable sintering process, the maximum density achieved was approx.

Actually, the smallest layer is 50 µ m.Of all the parameter combinations that lead to a stable sintering process, the maximum density achieved was approx. 50 % of theoretical density so that two direct applications arise: the laser sintering of casting components and the laser sintering of master patterns for metal spraying.Investment Casting ApplicationThe conventional investment casting process begins with the fabrication of wax patternswith the same basic geometrical shape as the finished cast part. These master patterns are normally made by injection molding. Once a wax pattern is produced, it is assembled with other wax components to form a metal delivery system, called the gate and runner system. The entire wax assembly is then dipped in a ceramic slurry, covered with sand and allowed to dry. The dipping, sanding and drying process is repeated until a shell of approx. 6 mm to 8 mm is applied. These process steps generally require several days.

The dipping, sanding and drying process is repeated until a shell of approx. 6 mm to 8 mm is applied. These process steps generally require several days. Once the ceramic shell is finished, the entire assembly is placed in a steam autoclave to remove the wax. Finally, the mould is preheated to a specific temperature and filled with molten metal, creating the metal casting [6].The major impact Rapid Prototyping processes have had on investment casting is their ability to make master patterns without the cost and lead times associated with fabricating injection molds, Fig.

2.449Conventional investment casting Conventional investment castingPattern making Pattern makingSlurring, sanding & drying of shell Slurring, sanding & drying of shellcasting castingfinishing finishingConventional investment casting using RP master patterns Conventional investment casting using RP master patternspattern pattern making makingSlurring, sanding & drying of shell Slurring, sanding & drying of shellcasting castingfinishing finishingtime savings time savingsInvestment casting using laser sintered shells Investment casting using laser sintered shellsShell making Shell making by SLS by SLScasting castingfinishing finishingtime savings time savingsFig. 2: Comparison of investment casting processesA direct generation of casting shells without the need of master patterns could have the highest time savings because the iterative process steps of slurrying, sanding and drying of master patterns are eliminated.

But since complete shells are generated the running systems and risers have to be designed and incorporated within the part geometry. A cavity is generated within the complete system with an external off-set of 6 mm to 8 mm (corresponding to the wall thickness of the shell).Before casting, the loose powder in the shell has to be removed and the shell cleaned. As usual, the laser sintered shell is preheated in an oven directly before casting. Fig. 3 and Fig. 4 show some examples of metal prototypes cast in laser sintered shells:Fig. 3: Impeller wheel cast in AlSi7450Fig. 4: Whirl chamber of a diesel engine cast in X15CrNiSi25 20The processing time for three casting shells of the impeller wheel respectively four shellsof the whirl chamber in parallel was only 10 hours, each. The dimensional accuracy of the cast prototypes is approx. +/- 0.6 % which is typical for investment casting.

The dimensional accuracy of the cast prototypes is approx. +/- 0.6 % which is typical for investment casting. The surface roughness of cast prototypes is worse (Ra ~ 12 µm) compared to conventional investment casting (Ra < 6.3 µm) so that a mechanical finishing/polishing may be necessary.Another problem of laser sintered shells is the lacking possibility to examine the qualityof inner surfaces and contours before casting with the exception of Computer Tomography (CT) or endoscopy.A comparable application of ceramic laser sintering is the use for investment casting cores. They are necessary for conventional investment casting if the metal part has complex inner geometries. Today, the core production is expensive and/or time-consuming since a metal pressing die or a mold for slip casting is needed. Thus, they are mostly used for mass production runs and not for prototyping.

Thus, they are mostly used for mass production runs and not for prototyping. The production by laser sintering could be a fast and favorable way especially for prototypes, single part productions or small batches. A typical core geometry is shown in Fig. 5:Fig. 5: Laser sintered water jacket core of a single cylinder (ZrSiO4)451Since the laser sintering parameters are not changeable to ensure the tightness of cores, the surface quality is still bad. But a dipping of the laser sintered core into a ceramic slurry increases the surface quality (Ra ~ 4.0 µm) so that laser sintered cores are also usable for production runs, Fig. 6.Fig. 6: Cast part split in two with an inner contour made by a laser sintered coreSpray Metal Tooling ApplicationMetal spraying belongs to the indirect Rapid Tooling where a master pattern is producedby Rapid Prototyping techniques. Afterwards, a metal alloy is sprayed onto the surface.

Afterwards, a metal alloy is sprayed onto the surface. Finally, the master pattern is destroyed, the mold insert is trimmed to the desired dimension and attached to the mother mold.This process is well-known for low-melting alloys (e.g. Kirksite or Zamac) which are sprayed onto a plastic master pattern. Compared to conventional mould making, prototype moulds could be available in better time. In 1999, Ford Motor Company licensed this technology for carbon steel [3, 7].Due to the high temperatures involved, ceramic master patterns are necessary in order tobe able to spray steel. The requirements on ceramic materials for steel spraying are [7, 8]:1. A low coefficient of thermal expansion,2. A high thermal shock resistance,3. A good compatibility to sprayed carbon steel and4. A fast processing with laser sintering.Ceramic powders like silicon oxide, silicon carbide, aluminum titanate and zirconiumsilicate are compatible with these four requirements, Fig. 7.452Fig.

A fast processing with laser sintering.Ceramic powders like silicon oxide, silicon carbide, aluminum titanate and zirconiumsilicate are compatible with these four requirements, Fig. 7.452Fig. 7: Sample parts (from upper left to lower right: SiO2, ZrSiO4, Al2TiO5 and SiC)While the laser sintering of silicon oxide led to an inaccuracy of the master pattern due tosporadic delamination the silicon carbide respectively aluminum titanate patterns were to fragile to handle. An infiltration with epoxy increased the strength but then the ceramic pattern were not removable by sand blasting so that zirconium silicate was used for practical tests [8].The following picture shows four laser sintered master patterns of a sheet metal formingdie, Fig. 8. The segments were produced on a laser sintering system with a maximum laser power of 200 W and a maximum work area of 250 mm by 250 mm (~ 9.8” by 9.8”).Fig.

8. The segments were produced on a laser sintering system with a maximum laser power of 200 W and a maximum work area of 250 mm by 250 mm (~ 9.8” by 9.8”).Fig. 8: Laser sintered ceramic master patterns453To prevent a relative movement of the segments during spraying, tongues and grooveswere designed. The base of each segment was approx. 245 mm by 245 mm. The average manufacturing time for each segment was approx. 30 hours. Without tongues and grooves the manufacture would be reduced by 50 %. After laser sintering the segments were compared with the 3D-CAD data. The average dimensional accuracy was better than 0.2 mm, a very good result and acceptable for sheet metal forming.Figure 9 illustrates the sprayed steel mold. The average surface roughness (Ra) was approx. 17 µm which is high compared to milled surfaces.

The average surface roughness (Ra) was approx. 17 µm which is high compared to milled surfaces. In addition, cross lines due to the fitting of the four segments were visible so that a post-machining and polishing would be necessary.Fig.8: Sprayed carbon steel moldThe surface quality could be improved by painting the surface with a ceramic slurrysimilar to the one used to produce investment casting cores. The post processing would be minimized thereby. This could form the subject of the next investigation leading to a faster production of steel molds.Conclusion and OutlookIt has been shown that ceramic laser sintering is a promising method of producing metal prototypes or moulds very rapidly. The examples shown illustrate that both ferrous and non- ferrous alloys can be successfully cast using laser sintered shells.

The examples shown illustrate that both ferrous and non- ferrous alloys can be successfully cast using laser sintered shells. The dimensional accuracy of the metal prototypes compare favorably with conventional cast parts.The advantages of laser sintered investment casting cores are that they are usable for complex prototypes as well as for single part productions since a slurrying increases the surface quality. A quality analysis could be done directly after laser sintering and not after casting. Finally, laser sintered cores are easier to remove after casting than conventional cores.454The use for spray metal tooling of stamping dies or sheet metal forming moulds is only limited by the present work area of laser sintering systems.

rect Laser Sintering of MetalsWilliam T. Carter, Jr. and Marshall G. Jones General Electric Corporate Research & Developlnent CenterAbstractThe use of a directed laser bealn source to selectively sinter multiple layers of binderless metal powder for the purposes of rapid prototyping is described. The work in this paper is restricted to -325 mesh iron powder, which was sintered using a C\V 50 W Nd:YAG laser to approximately 3.5% density. A subsequent post-treatlnent was perfornled to achieve a fully dense saulple. It is envisioned that such a system can be used to manufacture functional metallic prototypes directly from CAD without part-specific tooling.1Background1.1 Rapid PrototypingThe use of rapid prototyping in a concurrent engineering environn1ent results in re duced product developll1ent cycle tilne-a natural consequence of the quick availability of working testable hardware.

A functional prototype allows the identification of de sign deficiencies or areas for design inlprovell1ent; it allows experilnental stress analy sis, vibrations testing, and other design perfonnance tests with a very low lead-time. In some cases, experilnental analysis of prototypes has proven to be less expensive than cOlnputer-based analysis techniques such as finite elell1ent analysis [1]. A COln mon characteristic of rapid prototyping methods is that no part-specific tooling, such as a mold or die, is required to ll1ake the part. Several reviews of rapid prototyping are available in the literature [2] [3] [4].Use of rapid prototyping technology is expected to grow as Inore ll1aterials can be processed using the techniques. Many systell1s are cOlnlnercially available; the most popular technique to date is stereolithographYJ which generates a part from a bath of laser-cured photopolYll1er resin.

Many systell1s are cOlnlnercially available; the most popular technique to date is stereolithographYJ which generates a part from a bath of laser-cured photopolYll1er resin. Selective laser sirdering, the topic of this paper, is gaining popularity; it generates a part froll1 ll1ultiple layers of powder. COll1mercial systems for SLS are currently lilnited to ABS plastic, wax, nylon and polycarbonate materials.51To date, no c0111mercial freefor111 fabrication systeln is available for 111etallic ma terials. The preliminary results of an effort to address this issue are described in this paper.1.2Selective Laser Sintering (SLS)SLS generates a part by selectively bonding Inultiple layers of powder to build a three dimensional part in a layer-by-layer manner. The process has been comlnercially applied to ABS plastic, nylon, polycarbonate, and wax[5][6][7]. The process is very appealing for applications involving 111etals and ceranlics.

The process has been comlnercially applied to ABS plastic, nylon, polycarbonate, and wax[5][6][7]. The process is very appealing for applications involving 111etals and ceranlics. It is anticipated that SLS of metals will be less expensive and less ti111e-consu111ing than conventional production methods for 111etals, such as sand casting or investlnent casting, when the number of parts required is reasonably small. Therefore, the system will be used by designers and engineers during the prototyping stages of product developlnent. A c0111plete SLS apparatus will allow designers to quickly generate three-dilnensional functional prototypes of various parts. The process is also attractive to builders of custom tooling, dies and molds since these industries represent one-of-a-kind applications.Though there are no com111ercial systelns for applying this technique to metals, research is under way at the University of Texas at Austin.

1 One of their techniques is to coat the metal particles with a binder; the binder is selectively cured with a laser, and the part is later fired to burn out the binder and densify the part. Parts of final densities of 48% [8] and higher have been reported. Other work is ongoing to directly sinter Inetals by using a liquid-phase n1etallic material to fill the voids in a solid-phase powder, such as copper in nickel. Higher densities (82%) have been reported using this technique [9].The approach reported here differs in that the rnetal powder is sintered directly, without the introduction of a binder or a low-melting-te111perature liquid phase. A two-step process is proposed: The laser will not be used to generate a fully-dense final part; instead, the laser apparatus will be used to generate a "green" part that requires subsequent processing such as Jlot Isostat£c Press£ng (HIPing), oven firing, or infiltration.

The goal of this effort is to assess the process by laser sintering a sitnple cubic sample of iron, perfor111ing a HIP densification treat111ent, and evaluating the density of the resulting 111aterial.2 Experimental ApparatusA laboratory scale syste111 was constructed for evaluation of the SLS process. The system is capable of 111aking parts up to 1 in. by 1 in. by 3 in. using laser heating.lThe SLS process is patented by the lJniversity of Texas in Austin and has been reduced topractice by DTM, a subsidiary of B.F. Goodrich.52HopperMetering Plate Actuator PinRack and Pinion Shuttle Drive~ LJ..I..1..I..L.J..J..U..1..I..L.J..J..U..J...J..J...I...l.J.J..ll..Ii,.L1IU11J.l111J.,ijIl""1l.1- ••~1)~1!tJ9Adjustable :==\.

Metering Plate ,,,, ,,,,,,,Working Section \I~~;====="I II ITransport Chamber j j:II I IIi!Ii:'1 / i 11 1:----1:Waste BinShuttle Stepper MotorElevator Stepper MotorI I I I I I I I I I I I I I I I I I I I I I I I L_...JIIFigure 1: Powder handling system.2.1Powder Handling SystemThe powder handling system is shown schematically in Figure 1. This system includes two stepper motors: one to control the height of an elevator under the working section, another to control the location of a powder shuttle. To add a layer of powder, the elevator is lowered the desired distance, and the shuttle is moved until the actuator opens a powder metering device. A measured amount of powder is dropped through the transfer plate and into the shuttle transport chamber. The shuttle then returns to its home position. As the shuttle returns, a layer of powder is spread over the working section. Excess powder is dumped into the waste bin.

The 1.06 Jim wavelength laser bealu is channeled through a 600 Jirn single core step index quartz fiber optic to the powder handling systel11. At the terrninal end of the fiber optic, an output coupler and a lens focus the bean1 to approxiluately 0.5 lUlU dialueter on the luetal powder. The Nd:YAG laser is the better laser choice when cOlupared with a CO 2 laser because the 1.06 Jil11 heat source couples with lnetals luore efficiently than a 10.6 Ji111 heat source. See Figure 2. The Nd:YAG laser has the further advantage that, at higher powers, the bean1 is efficiently transluitted through fiber optics. This allows for more flexible SLS lllachine design.The output coupler asselubly is l1lounted on' an x-y table, which is translated via two additional stepper 1110tors.

This allows for more flexible SLS lllachine design.The output coupler asselubly is l1lounted on' an x-y table, which is translated via two additional stepper 1110tors. The laser's shutter and the four stepper motors (two in the powder handling systenl and two in the x-y table) are controlled using a single 80386 computer.2.3Secondary Processing (HIPing)The green form generated using the SLS systelu requires a post-treatluent to render a fully dense part. Since the part will, in general, be quite brittle and will possibly contain many small sections, sorrle Ineans of supporting the part during the treatment is required. Furthenuore, it is expected that the green structure will contain inter connected porosity that lllakes conventional gas HIPing impossible. The following technique was successfully iluplelnented to HIP laser-sinterecl specilnens:1. The speciluen was suspended in alu111ina powder in a Pyrex tube.2.

The following technique was successfully iluplelnented to HIP laser-sinterecl specilnens:1. The speciluen was suspended in alu111ina powder in a Pyrex tube.2. The tube was evacuated and baked at ~~ooo C for three hours to drive out all 1110isture and organic lllateriai. The tube was then sealed while still under vaCUU111.541.00.8~ -+-' > 0.6 -+-' U (J)'+- (J) cr:::0.411 (91 « I >-11 1 I I I - I0 1 l U~CarbOn1 1 1 C\ll0 U ISteel0.20.0110Wavelength (,urn)Figure 2: ~Wavelength dependence of reflectivity for various metals.3. The sealed tube was HIPed at 15 ksi and 11000 C for 15 minutes. The tem perature was linearly increased from room temperature at a rate of 50 0 C/min. and pressure was kept at 1 atm until a temperature of 9000 C was reached.4.

The tem perature was linearly increased from room temperature at a rate of 50 0 C/min. and pressure was kept at 1 atm until a temperature of 9000 C was reached.4. After cooling, the Pyrex tube was broken and the metal part was grit-blasted clean.3 Findings3.1 PropertiesA laser sintered specimen with a corresponding as-HIPed geometry is shown in Figure 3 along with the cross-sectional metallography of each part. This part was made from 44 Ilm (-324 mesh) iron powder in an environment of 10% hydrogen and argon using 7.5 W of laser power focused to a 0.5 mm spot. A scan speed of 5 in./min. was used with a scan spacing of 0.020 in. Each powder layer was approximately 0.004 in. thiclc The as-sintered part is approximately 35% dense, making it a brittle structure; however, it was able to withstand the small amount of handling necessary to place it55(3)(b)(c)Figure 3: Selective laser-sintered iron san1ple (a) before and after Inicrostructure.(b) and (c)in a HIP vessel.

The as- HIPed structure is very dense (> 99.9%) as shown in 3c, with only a small nU111ber of relnaining voids.3.2 EnvironmentThe most significant problerl1 associated with SLS of iron encountered in this this effort is the control of the sintering environn1ent and the forn1ation of oxide. reason, the powder handling systern was placed inside a glovebox and an overpressure of inert or fonning gas was applied inside the box. Additionally, the powder was "pre cleaned" by placing it in a fluidized bed charged with the san1e inert or forn1ing gas.563.2.1Inert Environn1entWhen using -325 mesh iron powder 1 it was found that a good single layer could be generated in an inert environn1ent of argon. UnfortunatelYl the bond between the layers was generally inadequate for the following reasons:~ The free energy of fonnation for oxide increases 'with decreasing ternperature [10].

UnfortunatelYl the bond between the layers was generally inadequate for the following reasons:~ The free energy of fonnation for oxide increases 'with decreasing ternperature [10]. Thus, in cooling to roon1 ten1perature fron1 a high homologous tempera ture, a period of high oxidation growth can be expected. This oxide is partic ularly undesirable in the SLS process because it prevents effective sintering or wetting.~ During the sintering of each layer, the ternperature of the powder was brought to a value above the nlelting ternperature of the n1etal. Each laser scan was perforn1ed relatively quickly. Thus, adjacent scans occulTed before the material cooled and sin1ultaneously oxidized. For this reason, the bonds within the layers (between scans) were acceptable.~ Between layers, the rnateria.l was allowed to return to rOOln ternperature.

For this reason, the bonds within the layers (between scans) were acceptable.~ Between layers, the rnateria.l was allowed to return to rOOln ternperature. Thus, each layer ran through a ternperature cycle that is very conducive to oxide (Oxide was observed in SENI EDAX probing.) formation as discussed above. Though the atmosphere was free of oxide 1 enough residual oxygen was present in the powder to fonn an oxide layer on the upper surface and prevent good bonds between the layers.3.2.2 Reducing AtlTIOsphereA corn1l10n reducing atmosphere for ferrous 111aterials is a mixture of hydrogen and nitrogen; this rnixture is often called "forl1ling gas." Laser sintering in this atmosphere proved to be unsatisfactory because of the fonnation of nitrides between each layer, preventing inter-layer bonds. (Nitriding is a C01111110n surface hardening technique for steels.

(Nitriding is a C01111110n surface hardening technique for steels. )When the same iron rnaterial was sintered in a reducing atrnosphere of 10% hydro gen and argon, excellent bonds between the layers were achieved. It should be noted that all other process pararneters (laser power density, scan spacing, layer thickness, etc.) were kept constant between the tests in argon, nitrogen and hydrogen, and those in argon and hydrogen.3.3 Post-TreatmentsFroln the materials standpoint, HIPing proved to be satisfactory for these preli1l1inary tests-it achieved a dense san1ple with a predicted reduction in the dilnensions of the specirnen. From the standpoint of FFF, HIPing is not attractive without a rnodel that gives a "process c01l1pensated" geornetry for the sintering operation. A Inodel of57the HIPing process is required to define an "as-sintered" part geolnetry.

A Inodel of57the HIPing process is required to define an "as-sintered" part geolnetry. This lnodel is not straight-forward because, for exanlple, holes and cavities which are telnporarily filled with ahllnina powder will densify at a different rate during HIPing than the adj acent lneta.l.3.4 WarpingBecause the heat is applied at the top surface, the top densifies to a greater extent than does the botto111 surface of each layer. Additionally, the upper surface cools froln a higher temperature than 111aterial below, causing lnore thermal contraction on the upper surface. for these reasons, each layer of sintered powder tends to warp upward, which is undesirable froln a powder-handling standpoint because the thickness of each layer is not unifonn. The solution adopted for this work was to build an "anchor" of thin sinteredlayers onto which the actual structure was constructed.

The solution adopted for this work was to build an "anchor" of thin sinteredlayers onto which the actual structure was constructed. This thin fralne gives SOIlle stiffness to the overall structure and avoids warping in lnany situations.Various other techniques are possi ble:Heating the powder bed. This 111ethod has been iIllplell1ented [11] for laser sintering of ll1etals with good success; it reduces the thennal strain by reducing the telnperature excursion. This ll1ethod has the added benefit that it helps prevent oxide build-up between layers."Knitting" the structure in such a way that the loose powder is allowed to flow to fill the ga.ps left by the densified ll1aterial.Bonding the prototype to a rigid sall1ple of the sall1e ll1ateria1 during the first stages of the sintering process, efFectively building an anchor.

This may be acceptable for SOine applications where the anchor can becOlne an integral part of the laser-sintered structure.ConclusionsThe proof-of-concept goal that still1ulated this work effort was successfully achieved. Binderless iron powder \vas sintered into a cubic shape of 35% density in a reducing atInosphere using a directed laser bearn. The cube was later HIPed to full density with a predicted reduction in overall din1ensions. It \vas dernonstrated that the control of the sintering environll1ent is critical for successful laser sintering. In this particular case, an environll1ent of 10% hydrogen and argon yielded the best inter-layer bonds.5 AcknowledgmentsThe work and advice of several people is acknowledged: Leon Beha, Jerry Harrison, Jim Resue, Steve Rutkowski, Paul Siell1ers, COSiIl10 l\10ntanaro, andl\1ark Brun.58References[1] G.W.

Royal, " Application of Stereolithography to Gas Turbine Design," Proceedings of the National Conference on Rapid Prototyping, June, 1990.[2] S. Ashley, "Rapid Prototyping Systems," lYlechanical Engineering, April 1991,pp. 36-43.[3] B.A. Beckert, "Cutting it in Rapid Prototyping," Cornputer>-Aided Engineering,Sept. 1991, pp. 28-40.[4] J.P. Kruth, "Material InCl'ess Manufacturing by Rapid Prototyping Techniques,"Annals of the CIRP, 40-2, 1991, pp. 603-614.[5] ILL. Marcus, et al., "Solid Freefonn Fabrication: Powder Processing," CeramicsBulletin, 69-6, 1990, pp. 1030-1031.[6] H.L. lVlarcus, et al., "Fro111 Cornputer to COInponent in 1.5 Minutes: The In tegrated l\1anufacture of Three-Dinlensional Objects," Joun1,al of ]\!!etals, April 1990, pp. 8-10.[7]Nutt, "The Selective Laser Sintering Process," Photonics Spect'ra, Sept. 1991,pp. 102-104.[8]Badrinarayan and J.Vv.

The combination of direct laser sintering and infiltration with metal melts allows the production of strong near net shaped components without shrinkage. A composite metal powder consisting of Ni, Cu, Sn and P was successfully sintered in a Selective-Laser Sintering unit in ambient atmosphere at room temperature. The influence of laser intensity on microstructure and sintering behaviour is discussed. Infiltration experiments were done with partially sintered samples. Full density could be achieved without shrinkage. Mechanical properties and microstructural development will be discussed.IntroductionCurrently laser sintering of metal and ceramic powders to manufacture parts is in research. Indirect sintering of polymer coated powders, where the polymer acts as a low melting point binder phase, has been developed to produce metal and ceramic components. Subsequent debinding and sintering is necessary to increase strength /1/.

Subsequent debinding and sintering is necessary to increase strength /1/. Direct metal sintering mostly involves the formation of a liquid through a metal phase with low melting point, e.g. Ni/Sn mixture, or by supersolidus sintering of a two phase, e.g. Cu/Sn-alloys /2/. Sintering of high temperature materials without binder phase needs high sintering temperatures. A high temperature laser sintering process (HTSLS) is under development /3/. Porous components made by SLS need further densification, either by sintering or by infiltrating the component. Further densification by sintering produces shrinkage of the cOlnponent. Infiltration with liquid metal seems to be more promising, however needs development of the technique for different powder systems. Porous94SiC bodies produced by SLS of coated powder particles have been infiltrated with liquid AI /4/. Also the infiltration of Fe-based porous components with copper alloys is very common /5/.

Porous94SiC bodies produced by SLS of coated powder particles have been infiltrated with liquid AI /4/. Also the infiltration of Fe-based porous components with copper alloys is very common /5/. This paper reports on research work in laser sintering of a multiphase powder mixture in a prototype version of a laser sintering unit. In order to increase density and mechanical properties sintered samples were infiltrated with liquid metal.ExperimentalThe powder material used in the experiments contains three different components, Ni and compounds of Cu, Sn and P. A micrograph of the powder is shown in Fig. 1. Laser sintering of test layers was done at ambient atmosphere with a 100 W C02 (CW) laser at different laser intensities and a beam scan speed of 200-800 mmls. Complex parts were built with layer thicknesses of 200 flm. The layers were investigated by SE and optical microscopy.

Complex parts were built with layer thicknesses of 200 flm. The layers were investigated by SE and optical microscopy. Bulk samples were sintered by laser sintering and for infiltration experiments by conventional sintering in Argon atmosphere. Infiltration of the porous samples was carried out in air at 500 aC, using a PbAg brazing alloy. Different infiltration times were used in order to achieve full density and improved mechanical properties. Tensile strength and hardness were measured from the as sintered and infiltrated samples.Results and DiscussionsSinteringThe thickness of the laser sintered layer was about 100 flm. Single layers were strong enough to be handled without damage. With increasing intensity of the laser beam the powder material goes from sintering to nearly completed melting. The microstructures of the sintered layers of 3 different laser intensities at the same scan rate are presented in Fig.

The microstructures of the sintered layers of 3 different laser intensities at the same scan rate are presented in Fig. 2 a-c and the corresponding cross sections of the individual layers in Fig. 3 a-c.low intensity: Sintering with low laser intensity results in partial melting of the binder phase. The layer is homogeneous over the thickness. Ni-particles, appearing bright in the SEM micrograph, are not wetted.medium intensity: At medium intensity wetting of all particles including Ni is observed however the top surface seems to be completely melted, while the lower part still shows individual particles.95SiC bodies produced by SLS of coated powder particles have been infiltrated with liquid Al /4/. Also the infiltration of Fe-based porous components with copper alloys is very common /5/. laser sintering of a multiphase powder This paper reports on research work mixture in a prototype version of a laser sintering unit.

laser sintering of a multiphase powder This paper reports on research work mixture in a prototype version of a laser sintering unit. In order to increase density and mechanical properties sintered samples were infiltrated with liquid metal.Experimentalexperiments contains three different components, The powder material used in Ni and compounds of Cu, Sn and P. A micrograph of the powder is shown in Fig. 1. Laser sintering of test layers was done at ambient atmosphere with a 100 W C02 (CW) laser at different laser intensities and a beam scan speed of 200-800 mm/s. Complex parts were built with layer thicknesses of 200 flm. The layers were investigated by SE and optical microscopy. Bulk samples were sintered by SLS and for infiltration experiments by conventional sintering in Argon atmosphere. Infiltration of the porous samples was carried out in air at 500°C, using a PbAg brazing alloy.

Infiltration of the porous samples was carried out in air at 500°C, using a PbAg brazing alloy. Different infiltration times were used in order to achieve full density and improved mechanical properties. Tensile strength and hardness were measured from the as sintered and infiltrated samples.Results and DiscussionsThe thickness of the laser sintered layer was about 100 flm. Single layers were strong enough to be handled without damage. With increasing intensity of the laser beam the powder material goes from sintering to nearly completed melting. The microstructures of the sintered layers of 3 different laser intensities at the same scan rate are presented in Fig. 2 a-c and the corresponding cross sections of· the individual layers in3 a-c.low intensity: Sintering with low laser intensity results layer is homogeneous over micrograph, are not wetted.partial melting of the binder phase. The thickness.

The thickness. Ni-particles, appearing bright in the SEMmedium intensity: At medium intensity wetting of all particles including Ni is observed however the top surface seems to be completely melted, while the lower part still shows individual particles.96high intensity: At high laser intensity the powder is nearly completely remelted at the top surface. Also the lower part of the layer is melted, however individual particles are still visible.Based on the results of the laser sintering experiments a complex shape of a small turbine, see Fig. 4, could be generated layer by layer. No curling and no distortion sintering. The component showed no shrinkage during laser occurred during sintering.conventional sintering: The conventionally sintered microstructure exhibits homogeneous wetting of all particles by the binder phase, as visible in Fig. 5. The material still has open porosity to allow further infiltration of the component.

5. The material still has open porosity to allow further infiltration of the component. The as sintered density is r-.J 70% of theoretical density.Low laser intensities result in a preferred microstructure in the as sintered state, remelting occurs only at medium and high intensities. Optimisation has to be achieved in improving the bonding at low intensities, for example by altering the atmosphere. Generally remelting should be avoided, because it will cause shrinkage and curling of the sintered layer. Also the microstructure is less homogeneous.InfiltrationComplete infiltration of the conventionally sintered components with liquid metal was achieved after 45 min total infiltration time. Fig. 6 shows the microstructure of the infiltrated sample. Partial reaction zones could be detected, however these are not generally visible in the material. Tensile strength of the 45 min. infiltrated material is not improved as compared to the sintered material but hardness is increased.

Tensile strength of the 45 min. infiltrated material is not improved as compared to the sintered material but hardness is increased. The reason for the unimproved strength lies in the presence of singular large pores in the infiltrated material, as could be observed in the fracture surface. These pores are filled very slowly, because capillary pressure is proportional to llr, where r is the radius of the capillary. Prolonged infiltration time leads to an increase in strength, also hardness is sligthly reduced.Table 1: Mechanical properties ofas sintered and infiltrated materialMaterial as sintered material 45 min infiltrated material 90 min infiltrated materialTensile Strength (MPa) 96 - 105 92 - 102 132Rockwell Hardness 25.2 89.6 69.597ConclusionsDirect laser sintering of metal powder mixtures needs careful control of laser intensities to avoid remelting of the powder material and to achieve a homogeneous microstructure of the sintered layer.

Using the described powder mixture direct sintering of more complex parts was done at room temperature. Infiltration of the sintered porous material is practical, however, property improvement is poor at the moment. Further improvement to achieve reasonable strength values of about 200 MPa tensile strength is necessary.References/1/ Badrinarayan B. and Barlow, 1. W., "Metal Parts from Selective Laser Sinter ing of Metal-Polymer Powders", Solid Freeform Fabrication Symposium, Austin 1992, 141-146/2/ Bourell, D.L., Marcus, H.L., Barlow, J.W., Beaman, 1.J., "Selective Laser Sintering of Metals and Ceramics", International Journal of Powder Metal lurgy (Oct. 1992) 28, (4), 369-381/3/J. McWilliams, C. Hysinger, J.1.

rect Manufacturing of Metal Prototypes and Prototype ToolsProf. Dr.-Ing. Fritz Klocke, Dipl.-Ing. H. Wirtz, Fraunhofer Institute of Production Technology, Aachen, FRGDipl.-Phys. W. Meiners, Fraunhofer Institute of Laser Technology, Aachen, FRGAbstractDue to the limited mechanical characteristics of the materials which can currently be processed using industrially available Rapid Prototyping machines, a lot of research is focused on the development of techniques which allow a direct manufacturing of metallic parts. These include Selective Laser Sintering and Controlled Metal Build Up. Both methods produce the workpiece not by removal of material but by a layerwise deposition and local melting or sintering of a powder material without part-specific tooling.Controlled Metal Build Up is a new Rapid Prototyping technique similar to Fused Deposition Modeling combining laser generating/welding with conventional 2Y2 dimensional milling.

Due to the excellent surface quality and high dimensional and form accuracy achievable this technology offers an interesting alternative to the with Controlled Metal Build Up, time consuming processes for the manufacture of prototype tools required for conventional limited production runs.With respect to Selective Laser Sintering, a test facility was developed for experimental investigations into the direct sintering of low and high melting metallic powders without the use of a polymer binder as well as ceramic powders.

Great potential is expected from Selective Laser Sintering concerning the prototyping of molds and dies.This paper will discuss current developments for these two techniques as well as point outpossible applications and future developments.IntroductionRecent developments in Rapid Prototyping have to a large extent been focused on the manufacture of prototypes with characteristics closely corresponding to the characteristics of the serial part being developed, allowing functional testing [1, 2, 3,4]. The production of a prototype suitable for functional testing requires the prototype to be identical to the serial part in shape, is being thoroughly adressed material and production process. While the first requirement the second requirement has only through e.g. constant sufficiently been met for a small number of plastics [1, 3, 4]. During 1995, two commercial Rapid Prototyping systems for the manufacture of metal parts (DTM RapidTool, EOS EOSINT M) were released [5, 6].

During 1995, two commercial Rapid Prototyping systems for the manufacture of metal parts (DTM RapidTool, EOS EOSINT M) were released [5, 6]. The fast generation of metallic tools e.g. for injection molding enablesimprovement of machine accuracy,141the production of prototypes in large quantities that exactly match the characteristics of the serial part.these two new Rapid Prototyping systems have some drawbacks. The RapidTool process brought into the market by DTM, utilizing a polymer-coated steel powder for Selective Laser Sintering, time consuming infiltrations and part shrinkage during burn-out of the binder [5]. EOS selectively sinters a bronze-nickel powder and parts also need to be infiltrated to achieve full density [6]. Both processes require extensive part finishing [5, 6].However,is characterized by severalthe IPT has developed a new process called Controlled Metal Build Up (CMB) in close cooperation with Albrecht Roders GmbH & Co.

KG, Soltau, Germany, which combines laser generating with high speed milling.To overcome the problems associated with part density and accuracy,Research being conducted in the field of Selective Laser Sintering is focused on themachinability of conventional tool steel to high densities as well as light metals and ceramics.Controlled Metal Build UpControlled Metal Build Up is based on laser generating, a technique pioneered at the Fraunhofer Institute of Production Technology IPT [7, 8, 9, 10]. A metal powder is blown into a focused laser beam and completely melted inside a nozzle, Figure 1.

A metal powder is blown into a focused laser beam and completely melted inside a nozzle, Figure 1. A fine bead of material is deposited beyond the nozzle onto the surface of the workpiece while the equipment is moved to scan the inner regions and the outline of the part.focussing opticlaser beampowder infeed via gaspowder nozzlelaser beamtrack producedtrackFigure 1:Powder nozzle used for CMB and principle of operation142Since neither the sharpness of the edges nor the levels of form and dimensional accuracy achieved meet the requirements of a pre-series tool, the contour of the part is milled. The upper surfaces also undergo a milling operation to keep the distance between the upper edge of the last deposited layer and the lower edge of the powder nozzle constant, Figure 2.

The upper surfaces also undergo a milling operation to keep the distance between the upper edge of the last deposited layer and the lower edge of the powder nozzle constant, Figure 2. Similar to the other Rapid Prototyping techniques, the workpiece is created layer by layer.process stepsFigure 2:Process steps of CMBCMB allows processing of numerous metallic materials ranging from bronze through steel to hard alloys which are frequently used to protect against wear. Since CMB is, in contrast to conventional cutting operations, a generative process, narrow, deep grooves may easily be produced using a constant, low engagement depth of the milling tool. The suitability of this method for automation in comparison to e.g. conventional 5-axis milling is a powerful advantage.

The suitability of this method for automation in comparison to e.g. conventional 5-axis milling is a powerful advantage. By virtue of the layer-by-Iayer nature of this technique, the CAD data may be processed quickly and with considerably less effort.the Fraunhofer IPT is essentially a high speed milling machine supplied by the Albrecht Roders company on which a laser generating head has been mounted, Figure 3. To date, simple geometries have been built using different materials in order to investigate the principle underlying the CMB process (Figure 3).The test facility atSelective Laser SinteringCurrent approaches to sinter metallic powders are concentrated on the direct or indirect sintering of the chosen material. Indirect sintering, deploying a binder phase (e.g. a polymer), has the advantage that only the binder material needs to be melted, most of the time at a low temperature allowing manufacturing of metallic parts using conventional plastic sintering machines.

However, the workpiece must be debindered and infiltrated, causing loss of accuracy and prolonging the time needed for manufacture. Direct sintering is aimed at melting the chosen material during the SLS process directly, eliminating the need for debindering. However, since processes currently industrially available leave the parts with a porosity of about 700/0, infiltration is still necessary to achieve full density. Another problem related to direct sintering of metallic143powders is shrinkage and the induction of thermal stresses resulting in distorted and curled parts. The development of materials with virtually no shrinkage (EOS) has to some extent elilninated this problem, however neglecting the desire to be able to produce parts out of the material used in series prod,uction.Figure 3:CMB test facility at the IPT and simple generated geometryA test facility was developed at the IPT for direct sintering of powder materials.

The process chamber of the test facility may be flooded with inert gas to avoid oxidization during sintering. The powder is supplied by a powder cylinder and spread across the part cylinder with an aluminum wiper. The test facility is equipped with aNd: YAG-Laser with a maximum output power of 300 W.Research conducted at the Fraunhofer Institute of Production Technology is focused on the direct sintering of metallic parts. Emphasis is put on achieving high densities during the SLS process by a careful adjustment of the process parameters, while maintaining a good surface quality. This way it is hoped that infiltration can be dispensed with without increasing the effort necessary for finishing.Figure 4 demonstrates the influence of laser power PL , scan velocity V ,I' and hatch spacing hI on typical workpiece quality parameters. Here, a single layer of bronze was sintered using 3 different energy densities by varying the above mentioned parameters according toE=h.l.

A possible explanation is that for small hatch spacings, most of the light of the laser beam is reflected or absorbed from the previously drawn, solid line which has a coefficient of absorption significantly less than the one of the powder. The amount of energy absorbed by the line melts the material again without adding further to the sintering depth. The amount of energy absorbed by the powder itself is small for small hatch spacings, thus not increasing the sintering depth any further.The density of the single layer increases with increasing energy density. A variation of hatch spacing or scan velocity has a greater effect on density than a variation of laser power, since the melted agglomerates of material are positioned closer together for slower scan speeds and smaller hatch spacings. Thus, an increase in scan velocity may not be compensated by an increase in laser power.The surface roughness increases with increasing laser power.

Thus, an increase in scan velocity may not be compensated by an increase in laser power.The surface roughness increases with increasing laser power. Higher laser powers lead to the formation of large, segregated agglomerates, being the reason for a poor surface quality. Decreasing the scan velocity also causes an aggregation of solidified material, increasing surface roughness until the velocity is so low that it allows formation of continuous solid lines. Once continuous solid lines may form, the surface quality is enhanced. Decreasing hatch spacing improves surface quality due to a continuous melting of small amounts of powder that attach to the previously solidified material.The Fraunhofer Institute of Production Technology is working in collaboration with the Fraunhofer Institute of Laser Technology on the development of SLS of steels, light alloys and ceramics.

By a careful adjustment of process parameters, especially scan vector length, the145Fraunhofer Institute of Laser Technology was able to produce SLS test parts made of stainless steel with a density greater than 950/0 and an acceptable surface quality. Figure 5 shows the test parts (the surface was milled to visualize the full density of the workpieces) and a cross-sectional view, displaying the low porosity_ A tool was sintered for verification of the suitability of this process for injection molding, Figure 6.Figure 5:Stainless steel test parts (lOx lOx 10 mm3) and cross-sectional viewFigure 6:Stainless steel tool for injection molding of rubber partsThe ILT and IPT were also able to sinter silicon carbide (SiC) directly without utilizing a binder material, Figure 7.

The resulting green part may be infiltrated with silicium and reaction sintered to a fully dense ceramic part.146Figure 7:Test part (10 x 10 x 3 mm3) sintered from SiC ceramicConclusionControlled Metal Build Up and Selective Laser Sintering are both promising methods for generating parts and tools out of powder materials. While two commercial SLS processes for the sintering of metal materials are commercially available, there are still many issues that need to be dealt with to make SLS of metals and ceramics applicable for widespread industrial use. Research at the Fraunhofer Institute of Production Technology is focused on the production of fully dense, accurate parts with a good surface quality through both of the above mentioned processes. The manufacture of various test parts have shown the feasibility of these methods for the rapid prototyping of parts and tools.

The manufacture of various test parts have shown the feasibility of these methods for the rapid prototyping of parts and tools. Current developments in collaboration with the Fraunhofer Institute of Laser Technology are aimed at gaining a deeper understanding of the pricinples of SLS of metals, especially steel, and extending the machinable range of materials to ceramICS.References[1]Venus, A. D. Opportunities for Selective Laser Sintering (SLS) in Integrated Rapid Product Development Proceedings of the 4th European Conference on Rapid Prototyping and Manufacturing, 13.-15.06.1995, Belgirate, Italy, pp. 55-67[2]Haferkamp, H., Bach, F.-W., Gerken, J., Marquering, M. Rapid Manufacturing - Direct Production of Metal Parts with Laser Radiation Proceedings of the 4th European Conference on Rapid Prototyping and Manufacturing, 13.-15.06.1995, Belgirate, Italy, pp. 123-135147[3]Forderhase, P., Ganninger, M., McAlea, K.

rect Metal Laser Fabrication of Cu slabs from powder precursor: surface depth of melt and furnace temperature issuesJorge Ramos-Grez1, Jorge Sanz-Guerrero, Tomás Larraín, Andrés RamírezMechanical and Metallurgical Engineering Department Pontificia Universidad Católica de Chile Av. Vicuña Mackenna 4860 – Macul – Santiago - CHILEABSTRACTA DMLF processing unit based on a raster-scanned 80 W CO2 laser beam has been developed to process single layers of metallic powder precursor. The process chamber provides atmosphere control (high vacuum and inert gas refill) besides temperature elevation up to 700 oC. In this work, copper powder precursor is confined inside a refractory steel mask surrounded by an aluminum oxide jacket that acts as insulator. The powder layers can have thicknesses of 0,5 and 1 mm. An infrared pyrometer measures in real time the temperature at one location in the surface of the powder bed.

The powder layers can have thicknesses of 0,5 and 1 mm. An infrared pyrometer measures in real time the temperature at one location in the surface of the powder bed. Scan speed, scan step, and furnace temperature have been varied to find combinations of such parameters that render surface melting and maximum densification. Partially melted samples were produced and their mass density was measured. Micro-hardness and surface roughness were also measured along the resolidified surface, the first rendering an average of 80,6 HV compared to the 90-105 HV of oxygen free copper, while the second resulting in an 8 μm Ra value.

Maximum melt of depth achieved is ~0,15 mm followed by a sintered layer.INTRODUCTIONThe “Direct Metal Laser Fabrication” (DMLF) process was developed at the end of the 90s by the inventors of the “Selective Laser Sintering” (SLSTM), which was patented and commercialized previously and successfully at the beginning of the same decade (Deckard 1989 [1] and Beaman et al 1997 [2]). The SLSTM process has allowed making a major step in the fabrication of complex geometry prototypes from polymers in freeform. This has also allowed shortening the design and fabrication cycle of high added value products. On the other hand, the DMLF process was developed afterwards with the objective to process metals in freeform directly, to recover high added value mechanical components (e.g., turbine blades) as well as obtain functional prototypes of complex geometry, this is, could withstand high temperatures and high mechanical stresses (Das et al 1998 [3]).

Need for a rapid manufacturing of metal molds and dies was another reason that pushed the development towards this type of technology (Atwood et al 1998 [4]; Lü et al 2001 [5]). However, it has to be mentioned that the DMLF process is one that still finds itself at a experimental phase and only the indirect-Selective-Laser-Sintering (iSLS, 3DSystems EEUU), Direct-Metal-Laser-Sintering (DMLS, EOS Germany) and Laser- Engineering-Net-Shaping (LENS, Optomec EEUU) have been commercialized in past years. Recent publications verified this argument (Venuvinod y Ma 2004 [6]; Erzincanh and Ermurat 2005 [7], Simchi et al 2003 [8,9]; Tang et al 2003 [10]; Zhu et al 20031 Corresponding author: jramos@ing.puc.cl246[11,12]; Morgan et al 2004 [13]; Ning et al 2006 [14]).

Among the main problems not been yet solved, we find: a resulting high superficial roughness of the component made this way (i.e., > 16 μm), presence of porosity and in certain cases a lack of densification (less that a 100%), and residual stresses (Nickel et al 1999 [15]). Oxygen and nitrogen contamination is yet another important issue that has to be considered, too.Most common metallic alloys used with the DMLF technology are those of Ti-Al-V (Wu and Mei 2003 [16]), Cu-Cu3P (Tang et al 2003 [10], Zhu et al 2003 [11, 12]), Fe-C-Cu- Mo-Ni (Simchi et al 2003 [8,9]) and Cu-SCuP (Ning et al 2006 [14]) among others. During 2003, Kyogoku et al [17] did preliminary trials using DMLF to process a Ti-Ni alloy. In this opportunity, Kyogoku et al showed the feasibility to produce polycrystalline structures with cylindrical shape (25 mm in length x 2-3 mm in diameter) by means of DMLF from a equi-atomic alloy composition of Ti and Ni.

These components showed the shape memory effect when analyzed by differential scanning calorimetry.During 2003, Ramos et al [18] carried out preliminary trial of DMLF at the University of Texas at Austin using Cu powder precursor. In that occasion, Cu specimens were obtained with the acceptable surface integrity. The grain size obtained varied from 10 to 100 μm in average transversal diameter. Longitudinally the grains showed a dendritic columnar morphology that was oriented towards the thermal gradient caused by the motion of the heat source. In this opportunity, room temperature conditions were kept at the interior of the processing chamber and a 240 W power laser was used at a wavelength of 1,06 μm.

In this opportunity, room temperature conditions were kept at the interior of the processing chamber and a 240 W power laser was used at a wavelength of 1,06 μm. The latter power level was enough to achieve complete melting of the pre- placed powder layer; moreover, at such wavelength enhanced coupling of the laser beam into the material takes place.THEORETICAL FRAME WORKThe Direct Metal Laser Fabricación (DMLF) process In the DMLF process, bonding of metal powder particles (i.e., 10 to 50 μm in diameter) is achieved by means of pressureless sintering at high temperatures and in other situations by the partial or total melting of a layer of powder rendering a liquid phase sintering process. In both cases, a high power infrared laser beam is used. The laser is raster scanned by means of an oscillating mirror system. The laser beam interacts over the surface of the powder layer, which has been previously compacted to thickness of 0,1 mm to 0,5 mm.

The laser beam interacts over the surface of the powder layer, which has been previously compacted to thickness of 0,1 mm to 0,5 mm. This layer is previously heated to a temperature in between 0,6 to 0,8 of the melting temperature of the powder, which is immerse inside an inert and reducing (e.g., Ar-H2). The energy provided by the laser beam increases the atmosphere temperature locally over the temperature at which sintering or melting take place (Beaman et al 1997 [1]). In this way, a bidimensional solid slab (solid layer) of low thickness is made.Reduction in the surface area of the powder is another driving force for the densification at ambient pressure (Bourell et al 2002 [19]; German 1984 [20]). Once a dense layer of material is fabricated this way, the process can be repeated many times rendering a multi- layered volumetric object as final result. The powder that surrounds the part and that has not been sintered or melted is afterwards removed manually.

The powder that surrounds the part and that has not been sintered or melted is afterwards removed manually. Among several other investigators, Zhu et al (2003) [21] have implemented a multi-layer DMLF process. (cid:2)247If the process is carried out using a single layer of compacted powder but of greater thickness (typically 1 mm to 4 mm), the result will also be a solid slab or rod with high degree of densification. However, under this configuration a phenomenon called “balling” takes place when the quantity of melted material surpasses a critical level, together with the effect of a surface tension gradient in the liquid and with the magnitude of the enthalpy of reaction between the liquid and the atmosphere above it. All these factors can induced surface ripples, and in some cases spherical agglomerates (Anthony and Cline 1977 [22]; Ramos 2003 [23]; Rombouts et al 2005 [24]).

This determines the temperature gradient that operates over the powder layer, determining the degree of sintering between layers or the fraction of melted material. Solidification of melted material is a function of the thermal gradient and the speed of the solidification front is related to the scanning speed of the laser beam. Time modulation of the laser beam during the melting regime also plays a critical role in densification of the resulting solid (Morgan et al 2004 [13]), however in this research this variable was not considered.Finally, it has to be mentioned that the large melting point of Cu, it affinity to become contaminated with oxygen and surface tension fluctuation in the liquid phase make this process very complex to control.

Due to the fact that the powder layer temperature can reach above 800 ºC, it is necessary to operate under vacuum conditions (i.e., higher than 10-2 torr) and then refill the furnace chamber with a reducing-inert gas to avoid residual oxygen contamination of the powders (Das et al 1998 [3]; Ramos 2003 [23]). To avoid grain growth in the specimens, due to the high temperatures at which the vacuum furnace is set, once the laser beam has been scanned, inert gas can be used to purge the atmosphere inside the furnace to further cool down the alloy fast enough to avoid structural transformation to take place. A virtual representation of the DMLF system developed and used in the present work is shown in Figure 2. (cid:2)248Scanning mirrorsBeam expansor- collimatorPyrometerLaserHigh vacuum and temperature furnace(cid:2)Figure 2.

(cid:2)248Scanning mirrorsBeam expansor- collimatorPyrometerLaserHigh vacuum and temperature furnace(cid:2)Figure 2. Virtual design of the available equipment to perform the Direct Metal Laser Fabrication process.Hypothesis The main hypothesis of the present work considered the feasibility to obtain surface melting and sintering of a powder bed of Cu powder by the DMLF process using a low power laser beam (80 W) if the right processing parameters combination is found: laser scan speed, percentage beam overlap and furnace temperature. A secondary hypothesis that evolved, is that at low laser power, initial temperature of the powder bed (i.e., temperature inside the furnace) beyond a certain value for a given powder precursor could hinder melt formation.EXPERIMENTAL SET UPIn July 2004 Ramos et al, started to develop a capacity to fabricate Cu alloys using the Direct-Metal-Laser-Fabrication (DMLF) technology with fundings form Fundación Andes in Chile.

This project contemplated setting up a system to do the DMLF process, integrating several subcomponents previously available as well as acquired with the funding available from that project (Figure 3). In this system, focusing is done using a F-theta lens with 36 cm of focal length resulting in a focusing diameter of 0,296 mm. Motion of the laser beam is achieved by a pair of scanning mirrors. Vacuum level and temperature inside the furnace are also operating up to 10-3 Torr and 700 °C, respectively. Figure 4, shows the pair of scanning mirrors located over the furnace window and the infrared pyrometer, which can measure surface temperature at one fixed location inside of the furnace. At present, the pyrometer is used to measure and control temperature at one fixed position by a PI closed loop scheme that actuates on the laser power intensity.

At present, the pyrometer is used to measure and control temperature at one fixed position by a PI closed loop scheme that actuates on the laser power intensity. The control systems behaves according to the following transfer function,G(s) =⋅ GK e 1 + 0,8s0,1s.The static gain, KG, has been determined empirically for a temperature range measured between 450°C to 900 °C at the surface of the copper powder, and its behavior is shown in Figure 5. It suggests that the heat source is no linear with respect to the gain and the percentage of power delivered. Nonetheless an average value of KG of 97,94454 was taken as a good approximation.249(1)Scanning MirrorsInfrared PyrometerHigh Vacuum Furnace(cid:2) Figure 3. DMLF system showing CO2 laser unit 80 W SYNRAD, an Edwards vacuum and temperature furnace, Cambridge Technology galvanometric mirrors.(cid:2)(cid:2)Figure 4.

This lack of laser power, narrows the experimentation range (leaving off materials systems with melting points above 1100 °C), so that in current research, laser power was kept at a maximum level. Moreover, the laser beam can perform a zig-zag or “snake” scanning pattern to simulate a linear heat source, as shown in Figure 6.250Figure 6. Laser beam raster scan patterns: zig-zag (up) and “snake” (down) type, showing different levels of beam overlap increasing from left to right.High temperature inside of the process chamber is achieved by an array of up to 6 loops of KanthalTM wire heating elements (2 to 3 mm in diameter) by a mechanism of heat radiation when passing a current of up to 75 amperes at a voltage of 18 volts (Figure 7). Maximum temperature achieved is 710 °C. The powder steel holder is located inside the volume surrounded by the heating element array.

Maximum temperature achieved is 710 °C. The powder steel holder is located inside the volume surrounded by the heating element array. It consists of a refractory steel mask having a 0,5 mm and 1,0 mm deep troughs and is surrounded by an aluminum oxide insulating jacket, as illustrated in Figure 8.Surface melted Cu specimenRegfractory steel crucibleFigure 7. Array of KanthalTM heating elements glowing inside of the processing chamber.Figure 8. Refractory steel mask and insulating alumina jacket used to hold the pre-placed Cu powder during the DMLF which has already been melted.EXPERIMENTAL TECHNIQUESSpecimen fabrication The copper powder precursor was preplaced inside the trough of the refractory steel mask, which was then inserted in the porous alumina enclosure. Masks had two different trough depths, 0,5 mm and 1 mm. The powder holder was then located inside of the processing chamber at the heating elements zone.

Masks had two different trough depths, 0,5 mm and 1 mm. The powder holder was then located inside of the processing chamber at the heating elements zone. The chamber is closed and251clamped and a mechanical pump draws vacuum up to 300 mTorr pressure. A coiled tube surrounds the chamber acting as a cooling water jacket. The chamber is then filled with argon gas, purged and then filled back again. The furnace heating elements are energized by increasing the amperage slowly up to 70 amps. This heating procedure takes ~20 minutes and then the laser beam is turned on to an 80 watts power level. The laser beam is expanded 8 times its size and collimated using a telescope, being then modulated in space by a pair of scanning mirrors. These reflect the beam downwards trough the F-theta lens which focuses it inside the chamber at the surface of the powder bed. The laser beam enters the chamber trough a ZnSe window.

These reflect the beam downwards trough the F-theta lens which focuses it inside the chamber at the surface of the powder bed. The laser beam enters the chamber trough a ZnSe window. Two process variations where considered, tuning off the furnace heating elements just before the laser beam started to scan and turning it off right after it finished scanning. Once the fabrication is accomplished, the specimens were allowed to cool down slowly for two hours until a temperature of 30 °C was reached inside the chamber.Characterization techniques An analytical weighting machine with four digit precision was used to weight the resulting specimens. Mass density was calculated by carefully measuring the dimension of the specimens by means of a micrometer to compute the volume. The infrared pyrometer was used to sample the temperature of the powder surface at one specific and fixed point (middle point of specimen width).

The infrared pyrometer was used to sample the temperature of the powder surface at one specific and fixed point (middle point of specimen width). A contact surface profile meter was used to measure arithmetic surface roughness of the specimens that experience melting, and a micro hardness Vickers tester was used to measure the surface micro hardness of those specimens, too.RESULTS AND DISCUSSIONThe characterization of the powder precursors as well as of the resulting specimens was done. Temperature histories at one fixed location in the powder bed surface were also measured. Results are presented as follows.Characterization of the Cu powder precursor The morphology of the Cu powder particles is shown in Figures 9 and 10 at different magnifications. Particle size ranged from 10 to 50 μm in diameter. These are not spherical particles but amorphous in shape. EDS chemical analysis was done on the powder to determine the presence of oxygen and other impurities content.

These are not spherical particles but amorphous in shape. EDS chemical analysis was done on the powder to determine the presence of oxygen and other impurities content. From Figure 11 presence of oxygen is negligible and the small aluminum peaks corresponds to the sample holder.Figure 9. Cu characterization, 1000x.powderprecursorSEMFigure 10. Cu powder precursor SEM characterization, 5000x.252Figure 11. EDS chemical analysis of the Cu powder used as precursor.Characterization of the laser processed Cu specimens A total of 18 specimens were fabricated, of which only 7 underwent melting at the surface of the sintered powder body. The parameters used to fabricate these specimens are listed in Table 1. All of the specimens where made with a laser power fixed of 70% it maximum capacity, approximately 66 watts before the F-theta lens. Figures 12 to 18 show macro pictures of the surface of all the specimens that melted.

Figures 12 to 18 show macro pictures of the surface of all the specimens that melted. Figure 19 shows a macro picture of Specimen #18, which did not undergo complete melting at the surface but instead liquid phase sintering. This specimen was however, fabricated under the highest average furnace temperature, almost 65°C more than Specimen #17. Another difference between these two specimens is that the latter was processed at a traverse speed twelve times faster.Table 1 Summary of processing parameters used in specimens that achieved surface melting (except for Specimen #18).

Figures 12, 16, 17 and 18 illustrate that for a 1 mm layer thickness, homogenously melted surfaces were obtained for Specimens # 3, # 7, # 12 and # 17 all along the length of the powder bed. The crater observed at the left side on each of these images, is due to the stationary position of the laser beam at the start of the fabrication and to the delay to reach its operating power level. Specimen #5 shown in Figure 14 was fabricated with a layer thickness of 6 mm contained in a ceramic crucible, surface melting was also homogeneously achieved, probably because the ceramic crucible retains heat more efficiently than the metal mask.Finally, from Figure 19, it can be seen that the surface of the Specimen # 18 has only melted partially favoring liquid sintering.

One hypothesis as to why it did not melted253completely, even though it experienced the highest average furnace temperature (725 °C) and the longer laser material interaction time (eight fold greater), is the fact that surface heat losses by radiation could have surpassed the heat input from the laser beam coupling. Heat losses by radiation heat transfer increase with the fourth power of the surface temperature according to Stefan Boltzmann´s Law. Moreover, the applied laser power in this process was not high (~55 W at the surface of the powder bed). Finally, according to Ujihara (1972) [25], reflectivity of metals is shown to decrease with increasing surface temperature, thus discarding the possibility of a surface reflectivity increase in the infrared region with increasing temperature.Figure 12. Specimen #3, laser power 66 W, Vs 20,1 mm/s, beam overlap 48%, Vt 0,94 mm/s, Tavg 499 °C, scan time 32 s. Layer thickness 1 mm.Figure 13.

Specimen #3, laser power 66 W, Vs 20,1 mm/s, beam overlap 48%, Vt 0,94 mm/s, Tavg 499 °C, scan time 32 s. Layer thickness 1 mm.Figure 13. Specimen #4, laser power 66 W, Vs 19,3 mm/s, beam overlap 48%, Vt 0,94 mm/s, Tavg 480 °C; scan time 33 s. Layer thickness 0,5 mm.Figure 14. Specimen #5, laser power 66 W, Vs 20 mm/s, beam overlap 48%, Vt 0,89 mm/s, Tavg 498 °C, scan time 32 s. Layer thickness 6 mm.Figure 15. Specimen #6, laser power 66 W, Vs 19 mm/s; beam overlap 48%, Vt 0,88 mm/s; scan time 34 s. Layer thickness 0,5 mm.Figure 16. Specimen #7, laser power 66 W, Vs 19 mm/s, beam overlap 48%, Vt 0,89 mm/s, Tavg 498°C, scan time 33 s. Layer thickness 1 mm.Figure 17. Specimen #12, laser power 66 W, Vs 22,7 mm/s, beam overlap 57%, Vt 0,92 mm/s, Tavg 568 °C, scan time 22 s. Layer thickness 1 mm.Figure 18. Specimen #17, laser power 66 W, Vs 17,8 mm/s, beam overlap 48%, Vt 1,2 mm/s, Tavg 660 °C, scan time 25 s. Layer thickness 1 mm.Figure 19.

Layer thickness 1 mm.Figure 18. Specimen #17, laser power 66 W, Vs 17,8 mm/s, beam overlap 48%, Vt 1,2 mm/s, Tavg 660 °C, scan time 25 s. Layer thickness 1 mm.Figure 19. Specimen #18, laser power 66 W, Vs 13,4 mm/s, beam overlap 95%, Vt 0,1 mm/s, Tavg 720 °C, scan time 205 s. Layer thickness 1 mm.Concerning the micro hardness Vickers measurements done on the specimens that showed complete surface melting, it can be seen in Table 2 that a value between 70 HV to 87 HV was recorded, which is lower when compared to the nominal hardness of an oxygen free copper sheet (90 to 105 HV [26]). Except for Specimen # 12, there is a trend between the measured mass density and the micro hardness HV at the surface of the specimens that melted as illustrated in Figure 20. On the other hand, the width W of the specimens can be twice as large as the scan pattern width programmed.

On the other hand, the width W of the specimens can be twice as large as the scan pattern width programmed. The surface arithmetic roughness (Ra) measured resulted near to 8 μm, which is lower than values reported in literature.254Table #2 Surface micro hardness Vickers, mass density and arithmetic surface roughness of specimens that experienced surface melting# 3 5 7 12 17B C1 C2 C A1 A2 A B1 B2 88 79 83 81 86 88 75 82 89 75 62 79 71 107 82 95 93 57 95 70 93 82 91 76 84 112 78 93 68 81 64 64 76 57 67 63 68 83 76 84 116 100 90 56 73Average HV 83 80 87 70 83ρ (kg/m3) 3083 2467 3348 3539 3073Ra W (μm) (mm) 3,45 NA NA 4,4 8 4,4 8 4 8 3,536003400) 3 m g k (/3200y t i s n e d s s a M3000280026002400798183858789Vickers Micro Hardness (HV)Figure 20 Plot of the mass density versus Vickers micro hardness of Specimens # 3, # 5, # 7 and # 17, showing a linear trend.Finally, Figures 21 to 24 shows several macro pictures of Specimen # 12.

The melted and resolidified surface shows a rippling pattern which could be due to surface tension gradients which tend to form surface waves that could freeze before their amplitude is completely dumped, moreover, pulling loose power into the melt front from the peripheral zones. The depth of melt in this specimen is ~0,15 mm followed by a sintered layer.Figure 21. Specimen # 12, melted front surface indicating rippling.Figure 22. Specimen # 12, sintered back surface.Figure 23 Specimen # 12, cross section indicating that the melted surface layer has a depth of 0,15 mm.Figure 24 Specimen # 12, closed up pf melted front surface indicating frozen surface waves.255Furnace temperature issues All specimens up to # 12 where fabricated so that the furnace heating elements were turned off when the laser beam started to scan over the powder bed surface.

For the remaining specimens, the furnace was turned off only after the laser beam had completely scanned the surface of the powder bed. Figure 25 shows a plot of depth of melt versus furnace temperature, resulting from a simplified lumped heat balance model (Equation 2), losses by radiation (including similar process parameters and constant temperature properties of copper taken from ref. 27 and ε = 0,9; η = 0,1; W = 3 mm; Α = 15 mm x W).

In the first plot, Figure 26, the change in temperature that takes places when the laser beam traverses the fixed spot, as measured by the pyrometer, is approximately 200 °C, for an average furnace temperature of 660°C. On the other hand, Figure 27 shows that an increase in temperature of the order of only 120 °C has taken place for a higher furnace temperature of 720 °C. Moreover, the relative increase in temperature of Specimen # 17 compared to that of Specimen # 18 is 20°C larger. Additionally, the latter specimen experience a much longer exposure time to the laser beam and that in both cases the laser beam power was kept constant.is counterit must be remembered that256(2)900)850C ° (e r u t a r e p m e t800750e c a f r u S700650200°C660°C6000510152025Time (s)Figure 26.

Surface temperature profile at one fixed location in Specimen # 17, laser power 55 W, Vs 17,8 mm/s, beam overlap 48%, Vt 1,2 mm/s, Tavg 660 °C.900850)C ° ( e r u t a r e p m e t e c a f r u S800750700120°C720°C650600050100150200Time (s)Figure 27. Surface temperature profile at a fixed location in Specimen #18 laser power 55 W, Vs 13,4 mm/s, beam overlap 95%, Vt 0,1 mm/s, Tavg 720 °C25730250CONCLUSIONSIt has been shown that it is feasible to do DMLF of preplaced copper precursor powder using a raster scan low power CO2 laser (66 watts before the F-theta lens) together with furnace temperature and inert gas atmosphere. Copper is known for its high surface reflectivity to infrared wavelengths, however, when in powder form its reflectivity apparently becomes smaller.In one hand, powder beds of 0,5 mm depth, melted completely but showed an irregular shape due to a lack of powder quantity.

On the other hand, due to the lack of laser power, it is not possible to melt the entire thickness of the 1 mm preplaced powder bed but only a surface layer of ~0,15 mm . However, this melt depth value is common to the multiple layer DMLF process. Full densification of the surface layer is achieved with a micro hardness HV value 20% lower than free oxygen copper. Overall mass density is almost a third of pure copper, due to the fact that below the molten surface a sintered layer has been produced with a 40% densification. Surface ripples are observed macroscopically, suggesting that surface tension gradient created surface waves in the liquid which then became frozen.A counter intuitive behavior has possibly occurred and been observed. Specimens fabricated under higher average furnace temperature appeared not to have melted completely at the surface while others with lower furnace temperature had, when under same laser power level.

Specimens fabricated under higher average furnace temperature appeared not to have melted completely at the surface while others with lower furnace temperature had, when under same laser power level. Temperature history profiles measured suggest that a larger temperature increase is obtained for lower furnace temperatures than at higher ones. A simplified lumped heat balance model clearly shows that for low laser power, there is a critical furnace temperature, above which depth of melts is reduced drastically as this temperature approaches the melting temperature, unless laser power is increased. Larger surface heat losses by radiation than heat intake from the laser beam could explain this unusual observation. More empirical data is required to validate it.A Cu-27wt%Ti alloy will be used a precursor material in the future, as it has an eutectic point at 820°C, almost 200°C lower than the melting point of pure Cu.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA DIRECT METAL LASER MELTING SYSTEM USING A CONTINUOUSLY ROTATING POWDER BEDWilliam T. Carter*, Michael E. Graham*, Christopher Hayden*, Younkoo Jeong*, Justin Mamrak†, Brian McCarthy*, William Monaghan*, Edward J. Nieters*, Victor Ostroverkhov*, Subhrajit Roychowdhury*, Andrea Schmitz*, Michael R. Tucker*GE Research, Niskayuna NY 12309†Concept Laser, A GE Additive Company, Lichtenfels, GermanyAbstractThe aircraft engine industry manufactures many metal parts of large diameter, but small cross-sectional area. Designers of these parts require increasingly complex geometries for improved aerodynamic efficiency and cooling. The combination of large diameter and complex geometric features inspired the development of a new Direct Metal Laser Melting (DMLM) architecture with a rotating powder bed.

The combination of large diameter and complex geometric features inspired the development of a new Direct Metal Laser Melting (DMLM) architecture with a rotating powder bed. The system coordinates the rotational motion of a powder bed with an ascending laser scanner and recoater to build in a helical fashion. A single-point powder feeder delivers metal powder near the inner radius of an annular build volume, and the recoater spreads the powder to the outer radius in a “snow plow” fashion. Because the recoater and laser scanner are installed at different angular positions, they operate independently and simultaneously. A prototype system was built to demonstrate this concept for an aircraft engine combustor liner (600-mm dia.

A prototype system was built to demonstrate this concept for an aircraft engine combustor liner (600-mm dia. x 150-mm ht.) and showed continuous laser utilization exceeding 97%.1IntroductionMultiple reviews of Additive Manufacturing technology [1-7] report that commercial Direct Metal Laser Melting (DMLM) machines of today share a common Cartesian architecture based on linear actuators arranged on three orthogonal axes. The part is built along one axis, the recoater motion is along another, and the gas flow is along the third axis (though it is sometimes parallel to the recoater motion). An inherent disadvantage with this arrangement is a requirement of sequential operation: the laser is idle during motion of the axes and the axes are idle during lasing.

An inherent disadvantage with this arrangement is a requirement of sequential operation: the laser is idle during motion of the axes and the axes are idle during lasing. Laser utilization, defined as the ratio of laser fusing time divided by the total build time, is an important metric in system productivity and investigators have devised various schemes to improve this ratio [8-11] while working within the Cartesian architecture. High laser utilization is difficult to achieve for large parts of small cross sectional area, such as thin-walled tubes, because lasing time is low and recoating time is relatively high.

High laser utilization is difficult to achieve for large parts of small cross sectional area, such as thin-walled tubes, because lasing time is low and recoating time is relatively high. Multiple lasers are required when the part dimensions exceed the field of view of a single scanner, a fact that further decreases laser utilization because laser exposure time is further reduced relative to recoating time.1665This paper discusses a DMLM system architecture based on a polar coordinate system where the rotational and vertical axes are increased monotonically throughout a build without reversal. The approach was first explored by researchers at the University of Liverpool and published in 2005 [12,13]. Powder dispensing, recoating, and lasing occur at different angular positions, allowing these operations to occur simultaneously and continuously, thus maximizing laser utilization as the part is built in a helical fashion.

Geometries well-suited to the polar arrangement are of large diameter and small cross-section. Parts meeting these characteristics are common in the aerospace industry where additive manufacturing is increasingly important as an enabler of performance improvements and cost reductions, achieved through innovative design and part-count reduction [14,15]. The cross-sectional area of a typical aircraft engine part (e.g., combustor liners, and engine casings) is approximately 1% - 5% of the powder bed area, so maximizing laser utilization is particularly important.2 The Rotary DMLM ConceptA schematic of the rotary system is shown in Figure 1 [16-19]. A build plate is mounted on a rotational stage, which is rotated at a rate that is synchronized with a laser scanner. As in conventional DMLM [20], the laser scanner directs rapid motion of a laser beam onto the layer of powder, causing it to melt and fuse to prior layers.

As in conventional DMLM [20], the laser scanner directs rapid motion of a laser beam onto the layer of powder, causing it to melt and fuse to prior layers. At an angular position that differs from that of the scanner, a powder feeder deposits powder and a recoater arm levels each layer of powder. The scanner, powder feeder, and recoater are elevated as the part grows in height allowing the part to be built in a helical fashion.The primary benefits of the new architecture include:Continuous operation of the laser scanner • Simultaneous powder recoating and lasing • Scalability to large parts of small cross section • Ability to build large parts with a single laser scanner • Precise dosing and economical use of powder • Local gas flow management • Straight-forward implementation of multiple laser scannersThe first four benefits in the list above play key roles in enabling high laser utilization and result in high build speed.

The last three items are added benefits over Cartesian machine geometries in terms of both process cost and product quality.The geometry of parts built on the rotating system need not be axisymmetric; the scan pattern is adjusted for any local features as the build plate is rotated. As will be described below, the speed of the rotating stage is continuously adjusted as the bed rotates to maximize laser utilization. Simultaneously, the powder feed rate is adjusted to minimize powder use. In an extreme case of non-axisymmetry, the rotational system can be used to build multiple individual parts that1666fit within the annulus between the inner to outer radii on the build plate. No restrictions on part geometry beyond standard DMLM restrictions apply to parts built in this manner.Figure 1. Rotational System Components.

No restrictions on part geometry beyond standard DMLM restrictions apply to parts built in this manner.Figure 1. Rotational System Components. The scanner, powder feeder, and recoater ascend as the part is rotated, building the part in a helical fashion.-----3Scan StrategySeveral scan strategies are under development. For the initial trial, the part to be built is divided into vertical layers and angular sectors as shown in Figure 2. Though not a requirement in the general case, each of the angular sectors spans an equal angle and the angle is chosen to allow an integer multiple of sectors per revolution; this ensures that all sector boundaries aligned on top of each other throughout the building of the part. This simplification will be abandoned in future work, where staggering of sector boundaries may be desirable.

This simplification will be abandoned in future work, where staggering of sector boundaries may be desirable. Scan path files defining mark and move operations for the laser scanner are generated for each of these sectors before building the part.Since the time requirement for scanning each sector is different, the rotation rate must be updated for each sector to maximize utilization. To accomplish this, each scan file includes an accurate estimate of the time required for completion of the sector [21,22]. A synchronization pulse is sent to the supervisory controller when a sector is completely within the view of the scanner.1667The required modification to the rotational speed is computed, and the rotary stage is appropriately accelerated or decelerated to ensure that the sector is completed before it leaves the field of view of the laser.

The scanner is controlled by a card that receives encoder feedback from both galvanometer positions and the rotary stage, so relative motion between the laser and powder are independent of the rotational speed. This approach works well as long as roundoff error in the sector timing calculation does not lead to a case where the sector cannot be completed in time. A small safety factor has shown to be satisfactory, leading to laser-on time less than 100%, but still greater than 97%. It is worth noting that synchronization between the scanner and rotation stage is maintained during all transitions of the rotational rate, so scanning is continued as the stage is accelerated or decelerated.Because the scanner is not centered above the build plate, the coordinate system of the scanner must be calibrated with respect to the coordinate system of the rotating part.

The calibration is specific to the machine geometry and configuration and, once applied, the scanner controller transforms the laser marking pattern to the specified location on the build plate as each sector passes through the scanner’s field of view [23].Recoater Blade (Powder is deposited nearthe inner diameter) ----~-----///Rotating Powder BedRotating PartScanner Field of View (Stationary)Sector BoundariesFigure 2. Scan Strategy. Powder dispensing and recoating are performed continuously as the build plate is rotated. Laser scanning of each sector starts when it is fully within the scanner field of view and the rotation rate is adjusted to minimize laser idle time between sectors.4System PrototypePhotographs of a prototype system that was built by the authors to demonstrate the rotary concept are shown in Figure 3. A laser-safe enclosure provides an inert gas environment for the process.

A laser-safe enclosure provides an inert gas environment for the process. A local gas plenum encloses the field of view of the laser and serves to sweep fumes out of the laser track and into a filtration system. (Many of the components of the gas plenum were 3D printed, allowing design creativity in the gas-handling system.) The photo shows a tall1668permanent inner wall inboard of the build plate; this wall prevents metal powder from pouring into the space at the inner diameter, saving substantial powder.Build Plate Powder Feeder Laser ScannerEnclosureFigure 3. System Prototype. The demonstrator system can build parts to 24" diameter and 10" height.4.1System ControlThe system control software is written in Visual C# and runs on a personal computer running Windows 7. Motion control of the rotational and vertical axes is done using an Aerotech A3200 device.

Motion control of the rotational and vertical axes is done using an Aerotech A3200 device. Analog and digital I/O is controlled with a GE RSTi-EP remote I/O system fitted with various analog and digital I/O modules. The consolidation power source is a single-mode fiber laser Atla Prime 700W from nLight, Inc. The scanner used for the system is a custom- integrated system based on Scanlab GmbH hardware. The motion system is connected to the control computer via a dedicated Firewire port. The IO module and the laser use an ethernet connection for communication. The scanner is controlled via an RTC5 control board that is connected to the computer backplane through PCIe.For modularity and scalability, the software is a multi-threaded architecture with each logical device having its own interface thread. An overall GUI thread takes user input, displays1669system information and controls system state including calibration, diagnostic and build.

An overall GUI thread takes user input, displays1669system information and controls system state including calibration, diagnostic and build. For post- build analysis, a low-priority “blackbox” thread periodically logs timestamped system conditions (e.g., temperature) as well as all changes to any hardware state (e.g., rotational velocity).4.2System PerformanceAn aircraft engine combustor liner (approximately 600-mm in diameter and 150-mm in height) was chosen as a demonstrator part (Figure 4). The part is thin-walled with a 20-mm thick flange and a 2-mm thick body. The part is complex, with a hole structure that includes large and small holes. Some of holes are too small to be seen in the photograph. The rotation rate was slowest (0.28 rpm) near the base of the part, where a thick flange required significant scanning time. The rate was fastest (2 rpm) at the mid-span where large horizontal holes were printed.

The rate was fastest (2 rpm) at the mid-span where large horizontal holes were printed. Sector-to-sector variations in build time required rotation rate changes with an average of 0.2%, but sometime reached 30% at positions with rotationally asymmetric features and as high as 113% at the part thickness transitional section. All accelerations were accommodated within the rotary stage sudden look-ahead algorithm was used capabilities. A acceleration/deceleration events to prevent missing a sector synchronization pulse. Laser utilization per sector varied between 75% and 99%, but the low-utilization excursions were brief resulting in an average utilization of 97.3% over the build duration (Figure 4). Though published data does not exist for conventional DMLM processing of a part of this size, the authors estimate that these systems would achieve approximately 30% utilization.

Though published data does not exist for conventional DMLM processing of a part of this size, the authors estimate that these systems would achieve approximately 30% utilization. Thus, the rotary system provides more than a 3x improvement in productivity.simpleto accommodate100%80%~C: 0 :;:. ~ 5 ~ .!!I60%40%20%- - Laser Utilization( %)A\lerage: 97.3% Powder Use (L)--------0%02040100 60 80 Build Height (mm)120140Figure 4. Combustor Liner. This demonstrator part was built using a CoCr Alloy. A single laser ran continuously with utilization per sector varying between 73% and 98%, averaging 97.3% over the entire build. Powder dispensed throughout the build totaled 27.6 L.The powder feed rate, also shown in Figure 4, was adjusted in coordination with the rotational speed of the part. The total dispensed volume of powder was 27.6 liters, with nearly half of the amount being due to powder feeding overflow outside the annular powder bed.

The total dispensed volume of powder was 27.6 liters, with nearly half of the amount being due to powder feeding overflow outside the annular powder bed. The volume of the consolidated part was 0.9 L, corresponding to 5% of total dispensed powder. For comparison, it would take 120 Liters of powder to fill a traditional rectangular powder bed for this part (including the typical overflow requirement). Thus the rotary system provides a 4x decrease in powder required.1670--1605040~ 30 ., :3 ~ 'O 3: 0 Q.201005 ConclusionsA rotating powder bed DMLM system based on polar coordinates was shown to offer significant advantages in build time and powder use over conventional Cartesian DMLM machines for many parts of interest to the aircraft engine industry. These parts are characterized as generally axisymmetric, of large diameter, and of thin cross section.

These parts are characterized as generally axisymmetric, of large diameter, and of thin cross section. The authors showed that laser utilization exceeding 97% was easily achieved while building a 600-mm diameter combustor liner with the new architecture. As added benefits, the system was able to build parts larger than the field of view of a single laser, and the metal powder requirement was reduced by a factor of 4.The authors are actively developing the rotary system to further increase productivity by adding lasers and increasing laser power. A larger system is under construction for building parts larger than 1-m in diameter and 1-m in height.6 References[1] Frazier, William E., “Metal Additive Manufacturing: A Review,” JMEPEG (2014) 23:1917- 1928[2] Kaufui V. Wong and Aldo Hernandez, “A Review of Additive Manufacturing,” Int.

rect Metal Laser Re-Melting (DMLR) of 316L Stainless Steel PowderPart 1: Analysis of Thin Wall StructuresRhys Morgan, Adam Papworth, Chris Sutcliffe, Pete Fox, Bill O’NeillResearch in Advanced Technologies GroupFaculty of Engineering, The University of Liverpool, UK.AbstractDirect Metal Laser Re-Melting is a process variant of Selective Laser Sintering, whereby 316Lstainless steel powder fractions are melted by a high power Nd:YAG laser. Layers are built up asa series of single lines to produce thin walled structures in the range » 0.3-1.0mm thick. Thestructures exhibit a periodic, angular roughness to the wall surfaces. The samples also display awave-like pattern on their upper surfaces. Further investigations reveal the angled ‘wave’ patternexists on a macroscopic level in the microstructure. The pattern is fully formed by the third orfourth layer. Fern-like grain structures follow the orientation of the surface roughness and existacross many layers.

The pattern is fully formed by the third orfourth layer. Fern-like grain structures follow the orientation of the surface roughness and existacross many layers. This is believed to be the effect of grain orientation within the samples. Themicrostructure reveals long, needle, cell structures. The uni-axial needles grow epitaxially fromprevious layers. The samples have been shown to exhibit very little or no thermally inducedresidual stresses.IntroductionSelective Laser Sintering has three main production routes for the processing of metalliccomponents. The first two, Indirect Metal Sintering and Binary Phase Sintering involve, at somestage in the process, the use of alloy materials with properties poorer than that of the highstrength, high temperature material in order to increase density or to improve melting and wettingcharacteristics (Tobin 1993; Bunnell, 1994).

The parts produced, however, will not reproduce theexact materials characteristics of a similarly machined component.The third route utilises only a single phase powder of the engineering material required for thefabricated end component. In this process, a high power laser beam directly melts and fuses thehigh temperature powder particles. There are, however, very significant problems associated withdirect illumination of engineering alloys in this way. Firstly, large thermal gradients exist in thelayers due to the non-homogenous heating of the powder by the laser. This causes substantialthermal stresses to be generated in the parts (Deckard, 1995). In polymer sintering processes,increasing bulk powder bed temperature reduces these thermal gradients.

This causes substantialthermal stresses to be generated in the parts (Deckard, 1995). In polymer sintering processes,increasing bulk powder bed temperature reduces these thermal gradients. However, due to thehigh temperatures involved with direct laser sintering, anchor plates have been used to overcome276the effect of thermal stresses in the laser-sintered part, and can be incorporated into thegeometry of the final component (Meiners, 1998).Final part density is a significant barrier to the success of single phase direct laser sintering. Themolten volume is shaped through surface energy minimisation, driven by surface tension. Thereis also a possible effect of free convective “Marangoni” forces, as experienced in melt poolsduring laser welding (Lampa, 1997). These effects act to shape the volume into strings ofconnected melt beads, which are inherently difficult to tessellate and therefore create areas ofporosity.

These effects act to shape the volume into strings ofconnected melt beads, which are inherently difficult to tessellate and therefore create areas ofporosity. Also, surface tension forces act to drag surrounding particles into the melt, causing localdepleted areas of material (Song, 1997). Furthermore, high temperature oxidation of the materialreduces wetting at the interface between layers and causes a greater degree of balling. Thesefactors must be overcome for the successful production of fully dense, engineering alloycomponents for heavy duty, functional use. Research at the University of Liverpool isinvestigating Direct Metal Laser Remelting (DMLR) of stainless steel with Nd:YAG laser toachieve this goal.DMLR Experimental ArrangementThe DMLR test facility consists of a Rofin Sinar 90W flash Q-Switched, Flash Lamp PumpedNd:YAG Industrial Laser Marker and an atmospheric control chamber built in-house at theUniversity of Liverpool.

It has a single powderfeed cylinder with counter rotating roller mechanism for powder deposition onto the build platform.Layer thickness is controlled by stepper motors, whose minimum linear step size is 1m m. Theprocess is controlled by in-house software for full automation and communication between thesystem and laser marker. The system is enclosed to allow for evacuation of ambient environmentand re-fill with Nitrogen process gas. No heaters are used before or during the process.The material under investigation is Stainless Steel 316L. The powder is gas atomised (spherical),specified as 80%<22m m, with a distribution approximately -<1m m < f < 60m m. It has thecomposition: 16.73Cr, 13.19Ni, 0.017C, 0.710Si, 2.69Mo, 1.6Mn, Fe balance (%wt).Samples were ground and polished to 1m m, and etched with standard austenitic (40% HCl, 20%HNO3, balance H2O) etch.

A further “coloured” etch was employed to establish variations incomposition withinthe austenite (4% Ammonium Hydrogen Difluoride, 16% PotassiumMetabisulfite, balance H2O). This etchant highlights microstructures, which are chromium rich inblue, while nickel rich regions are shown in brown.Experimental StudyInitial experiments were undertaken to assess the effect of laser and scanning variables on cold,loose powder beds of arbitrary depth. Detailed results can be seen elsewhere (Morgan, 2001). Asignificant effect of scanning on loose powder layers, was the significantly increased beamabsorption during the first scan lines, leading to increased volume of the coalesced melt beads,previously recognised as First Line Scan Balling (Hauser, 1999). The increased volume of thesefirst scan lines causes problems when applying further powder layers.

The increased volume of thesefirst scan lines causes problems when applying further powder layers. In order to overcome thiseffect, multiple layer experiments are undertaken on 3mm substrate plates of the same material.a) Wall produced at 80W, 50mms-,1, CW modeb) Wall produced at 80W, 70mms-,1, CW modec) Wall produced at 80W, 100mms-,1, CW modeFigure 2: Thin Wall Structures Produced at Varying Scan Speeds278Ten single scan -line walls, of length 15mm, were produced per plate. The layer thickness was aconstant 100m m. A hundred layers were produced during each experimental run. Due to theincreased conduction effects from the substrate, power was kept at a maximum to ensuresufficient fusion between the powder layer and the plate. Average power was measured at 80W. Scan Speeds were varied in the range 10 -500mms-1. For this study, the samples were produced using CW mode scanning.

Average power was measured at 80W. Scan Speeds were varied in the range 10 -500mms-1. For this study, the samples were produced using CW mode scanning. The beam was set at the focal position, with a diameter of 100m m.Results and DiscussionFigure 1 shows a typical plate with thin wall structures. Each wall is produced with a different scan speed, incremented by 10mms-1. Walls produced in the range 10-60mms-1 exhibit fulldensity, with a diagonal pattern emerging in the surface as shown in Figure 2a. A threshold speed of 70mms -1 creates a step in the diagonal pattern resulting in areas of porosity within the structures (Figure 2b). Above 90mms -1, the walls become a series of weakly connected verticalbranches; the thickness of each branch reducing with increasing scan speed (Figure 2c). At scan speeds above 140mms-1, there is insufficient energy imparted to the powder to affect bondingwith the underlying layer, resulting in failure of the build.

At scan speeds above 140mms-1, there is insufficient energy imparted to the powder to affect bondingwith the underlying layer, resulting in failure of the build. The thickness of the samples varies withscan speed. Increased thickness at the lower scan speeds is attributed to increased energyincident on the powder layer, resulting in radial conduction to the local surrounding powder,causing an increase in lateral melt volume and hence thicker walls. Also, there is a tendency forwall thickness to increase slightly with build height. This is due to the decreasing conduction2mm100m ma) Diagonal structure in polished sampleb) Cross Section of thin wall(4 layers)Figure 3: Formation of wave pattern in thin wall structures279effect from the substrate with increasing build height.

The samples show no sign of deformationor residual stress; this was borne out when samples were annealed at 500(cid:176) C for 4 hours andretained their form.1.5mm500m ma)’Fern’ Grain structure in thin wall sample b) Ni rich regions on upper surface of eachlayer (circled). Bulk of layer appears Cr richFigure 4: Analysis of microstructure with colour etchMacroscopic analysis of the polished samples reveals the pattern on a macro-scale within themicrostructure, while the upper surface of the samples exhibit a wave-like form (Figure 3a). Thewave pattern has a periodicity of approximately 1.5mm. The wave structure evolves from thesecond or third layer; by the fourth layer it is fully established.

Thewave pattern has a periodicity of approximately 1.5mm. The wave structure evolves from thesecond or third layer; by the fourth layer it is fully established. A cross sectional view of the firstthree layers shows that a possible cause of the structure is due to the melt volume falling over theside of the wall (Figure 3b).On closer examination of the polished samples, fern-like structures exist running along thediagonal pattern. These structures were initially thought to be the formation of nitrides. However,thermodynamically, this is not likely, as it is doubtful that there will be sufficient energy, in the timeavailable, to disassociate the nitrogen to form a nitride. Also, replacing the Nitrogen process gaswith Argon yielded the same grain structures.

Also, replacing the Nitrogen process gaswith Argon yielded the same grain structures. Using the colouring etch, the fern structures appearbrown surrounded by light blue regions, suggesting regions of nickel rich grains (Figure 4a).However, it is more likely that the ferns are due to grain orientation within the microstructure.Further analysis of the samples with the colouring etch reveals the upper surface of each scanline is rich in nickel compared with the chromium rich bulk material of each layer (Figure 4b).

Upon the passing of the beam, the underlying substrate (previouslayer) is re-melted, causing the cellular structure to exhibit uni-axial, epitaxial growth. Thecolumnar growth will continue throughout subsequent layers unless terminated by insufficientmelting of the boundary substrate (previous layer) as marked in Figure 5b.ConclusionsThe Direct Laser Re-Melting process has shown to produce full density thin walls (<1mm) at scan speeds < 70mms-1. The thin walls exhibit a wave-like pattern visible in the microstructure. Thispattern is believed to be due to melt dynamics. Within the wave pattern, there appears a fern-likegrain structure. This is believed to be due to grain orientation. The microstructure has shown fullmelting of the powder layers with no occurrence of unmelted particles. Grains exist along manylayers, confirming melt metallurgic bonding between layers. Surprisingly, there appear to be noresidual stresses in the walls.

rect Selective Laser Sintering and Containerless Hot Isostatic Pressing for High Performance Metal Components Suman Das, Martin Wohlert, Joseph J. Beaman, David L. Bourell Laboratory for Freeform Fabrication University of Texas at AustinAbstractA novel net shape manufacturing method known as SLSIHIP that combines the strengths of selective laser sintering (SLS) and hot isostatic pressing (HIP) is presented. Direct selective laser sintering is a rapid manufacturing technique that can produce high density metal components of complex geometry with an integral, gas impermeable skin. These components can then be directly post-processed to full density by containerless HIP. The advantages of in-situ HIP encapsulation include elimination of a secondary container material and associated container-powder interaction, reduced pre-processing time, a short HIP cycle and reduction in post-processing steps compared to HIP of canned parts.

Results of research conducted on Inconel 625 superalloy, Ti-6AI-4V and Monel are presented. This research is funded by DARPAlONR contract N00014-95-C 0139 titled "Low Cost Metal Processing Using SLSIHIP".INTRODUCTIONSelective laser sintering (SLS) is a layered manufacturing technique that can producefreeform three-dimensional objects directly from their CAD models without part specific tooling or human intervention. Parts are built by selectively fusing layers of apowder material using a scanning laser beam. Details on this process are described elsewhere1 technology for prototyping parts in a variety of polymeric materials and for creating investment casting patterns has been commercially developed by DTM Corporation.

More recently, RapidTool™ technology for creating prototype injection molding tooling has been introduced3 • This process is an indirect SLS technique that involves the use of polymer coated metal powders to produce a produce a green shape that is sUbse~uently post-processed by binder burnout and infiltration to produce a fully dense object4 prototypes at reduced cost and short lead time has spurred the development of so-called direct fabrication processes. A number of such next-generation direct fabrication processes are under development. These processes typically use a concentrated energy source to consolidate and produce a fully dense or nearly fully dense shape directly from constituent materials. The next generation of selective laser sintering, Le. direct fabrication of functional metal and cermet components and tooling is under development at the University of Texas 10,11,12.

The next generation of selective laser sintering, Le. direct fabrication of functional metal and cermet components and tooling is under development at the University of Texas 10,11,12. To produce full density metal parts having complex geometry, a novel net shape manufacturing technique called SLSIHIP is under development at the University of Texas. The idea is to consolidate the interior of a component to 80% or higher density and to fabricate an integral gas impermeable skin or "can" at the part boundary in-situ. The SLS processed part can then be directly post-processed by containerless HIP to full density. A final machining step will result in a part having the desired geometry and mechanical properties., • Selective laser sintering,5,6,7,8, • Demand for low volume functional metal81BACKGROUNDThe Department of Defense has a number of high value, high performance metalcomponents in service that are produced by hot isostatic pressing.

These parts are typically produced by conventional HIP of canned metal powders. Shaped metal cans are commonly used to encapsulate metal powders for HIP. The sheet metal container material is chosen so as to minimize interaction with the powder at processing temperatures. The container material must be removed by machining or by chemical methods after HIP post-processing. Properties of the can material including its melting temperature impose processing limits on the shaped metal encapsulation method.Complex shapes are typically produced using the ceramic mold process 13 developed by Crucible Materials. This process is similar to investment casting in that dry powder instead of molten metal is poured into a ceramic mold. The production of a near net shape is advantageous because it minimizes scrap losses and machining steps.

The production of a near net shape is advantageous because it minimizes scrap losses and machining steps. However, outgassing and heating cycles are long during this process because the ceramic mold is surrounded by a large volume of pressure transmitting medium14 ceramic mold method make it a time consuming and expensive process. In addition, non metallic contamination is also possible.The long cycle time and pre-processing steps necessary in theIn the SLSIHIP process (Figure 1), the component is produced by selectively consolidating a metal powder with a laser beam layer by layer. While producing each layer, a gas impermeable high density skin (> 98% density) is formed at the boundaries of the part. The interior of the part is laser processed to a high density typically exceeding 80%. Thus, the part is shaped and canned in-situ. The encapsulated part is evacuated, sealed and post-processed by containerless HIP to full density.

Thus, the part is shaped and canned in-situ. The encapsulated part is evacuated, sealed and post-processed by containerless HIP to full density. A final machining step may be applied if necessary.SLSIHIP has several advantages over conventional HIP methods. Since an integral skinor "can" is formed of the same material as the part, a secondary canning step is not necessary. The part is directly post-processed by containerless HIP. Adverse container-powder interactions are eliminated and post-HIP container removal is not required. SLSIHIP allows production of complex shapes at reduced cost and shorter lead-times.Based upon on a survey of several naval installations 15, four candidate materials havebeen selected for SLSIHIP process development. These are Inconel® 625, Ti-6AI-4V, 17-4PH stainless steel and Molybdenum.

These are Inconel® 625, Ti-6AI-4V, 17-4PH stainless steel and Molybdenum. Table 1 below summarizes demonstration components and their applications.ComponentMaterial Inconel625 Ti-6Al-4V 17-4PH SS Molybdenum Table 1 SLSIHIP demonstration components(Engine Vane lHousing base tink assembly Rotary valveApplicationlAircraft turbine engine component Tactical missile body component Tactical missile launcher component Torpedo component82Laser beam scans high density skin for current cross-sectionLaser beam scans interior of current cross-sectionIn-situ canning SLSIHIPDirect SLSFinal MachiningHot Isostatic PressingFigure 1 The SLSIHIP ProcessEXPERIMENTALCandidate powders for screening trials were obtained from Anval Corp. and NuclearMetals Inc. Anval provided Argon atomized Anval625 alloy powder (16-44 J.lIIl). Ti-6Al-4V (37-74 J.lIIl) produced by the PREP method was provided by Nuclear Metals Inc.

and NuclearMetals Inc. Anval provided Argon atomized Anval625 alloy powder (16-44 J.lIIl). Ti-6Al-4V (37-74 J.lIIl) produced by the PREP method was provided by Nuclear Metals Inc. The compositions of Anval 625 and PREP Ti-6Al-4V are shown in Table 2 and Table 3 respectively.CSiMnPSCrNi MoTiNbAlCoNFe0.028 0.160.04 0.007 0.008 21.362.28.220.273.360.210.04 0.051 4.00Table 2 Manufacturer's composition of Anval 62583AlVFe0CNHY6.354.190.190.190.020.01<0.01< 0.001Table 3 Manufacturer's composition of PREP Ti-6AI-4VCharacterization of the powders used in the SLSIHIP experiments included observationof the powder morphology by SEM and measurement of the powder surface area by gas adsorption surface area analysis (BET method).MorphologyBoth the Inconel625 and Ti-6Al-4V powders were quite spherical, as would be expectedfor gas atomized and PREP powders.

The high degree of sphericity provided excellent flow characteristics. A free flowing powder is important for SLS since the powder layers must be smooth and uniform. The presence of some particle rearrangement as material is drawn into the melt pool beneath the laser beam has also proven beneficial. The micrographs indicate a dendritic surface structure for both powders, but the presence of some segregation is not problematic since the material is melted and resolidified during the SLS process.Figure 2 SEM micrographs of -325 mesh Anval625 (left) and -200 mesh PREP Ti-6AI-4V (right)Surface Area AnalysisThe surface area of the powder is of greater concern, since the material must be thoroughly degassed prior to SLS processing. There are indications that the presence of oxide formation during SLS processing is a significant factor in influencing both the mechanical properties and the part density following the SLS stage.

There are indications that the presence of oxide formation during SLS processing is a significant factor in influencing both the mechanical properties and the part density following the SLS stage. A 5-point BET surface area analysis (Table 4) was performed on the powder using a Micromeritics ASAP 2010 Surface area Analyzer. Samples of Inconel 625 and Ti-6Al-4V weighing 26 gm and 18 gm respectively were degassed at 3500 C for 24 hours.84TiBallMaterial IInconel 625 IPREP Ti-6AI-4V Table 4 BET Surface area analysisPredicted Area (m2/f!.) 0.035 0.034Measured Area (m2/f!.) 0.052 0.0686Leak testingTo screen specimens produced by SLS for impermeability, a leak testing apparatus andprocedure was adapted from the Metals handbook article on containerless HIP I6 to be acceptable for HIP, a gas impermeable barrier must be provided. A helium leak rate less than 1 X 10-9 standard cm3/s is considered acceptable.

A helium leak rate less than 1 X 10-9 standard cm3/s is considered acceptable. Such low leak rates are required because the leak rate at typical a HIP pressure of 1000 atm (100 MPa) will be five orders of magnitude greater than that during leak testing at 1 atmosphere.For a specimenFeasibility demonstrationTo demonstrate feasibility of constructing a thin walled can acceptable for HIP, a rectangular geometry was chosen. The procedure established to qualify specimens and to refine SLS processing parameters is shown in the flowchart of Figure 3. A schematic of the test geometry is shown in Figure 4.SLSRefineMicrostructure EvaluationLeak testPassFailLaser scan patternHlPBuild 1direction_-.~.

A schematic of the test geometry is shown in Figure 4.SLSRefineMicrostructure EvaluationLeak testPassFailLaser scan patternHlPBuild 1direction_-.~. 1°4'Microstructure Evaluation410.125"1.2"Figure 3 Specimen qualification procedureFigure 4 Thin wall specimen schematicSLS apparatusSLS trials were conducted on a high temperature selective laser sintering machine designed and built at the University of Texas. This machine is equipped with a 250 Watt Nd:YAG laser, powder preheating capability up to 600° C and controlled atmosphere.85HIP apparatusHIP trials were conducted on a ABB model QIH-3 equipped with a graphite heating element. Temperatures up to 2000° C and pressures up to 200 MPa are attainable with the graphite element in inert atmosphere.

Temperatures up to 2000° C and pressures up to 200 MPa are attainable with the graphite element in inert atmosphere. Argon gas was used as the pressure transmitting medium.RESULTS AND DISCUSSIONTo demonstrate feasibility of constructing gas impermeable thin walls by SLS, specimensconforming to the geometry of Figure 4 were produced. These were then leak tested and screened according the procedure in Figure 3. Shown on the left in Figure 5 is the cross-section of a thin wall specimen produced by SLS. This specimen passed the leak test with a helium leak rate of less than 1 X 10 -10 standard cm3/s. This specimen was post-processed by a HIP cycle consisting of 1 hour at 1100 °C and 30 mTorr pressure followed by 45 minutes at 1100° C at 66 MPa. A cross-section of the HIPed specimen is shown on the right. Average porosity was measured over a metallographic montage in as SLS processed and HIPed specimens revealing a decrease in average porosity from 4.5% to 3.7%.,':.,..

Average porosity was measured over a metallographic montage in as SLS processed and HIPed specimens revealing a decrease in average porosity from 4.5% to 3.7%.,':.,. . .,',...-; ,. ",~ ,.r-.. t.j.. ... ,""" ~Figure 5 InconeI 625 Specimen 121496, as SLS processed (left) and HIP post-processed (right)The next step in SLSIHIP process development was to demonstrate capability ofconstructing a simple shape. A 0.5 inch diameter, 0.5 inch long cylinder was chosen for this purpose. Shown in Figure 6 is an axial cross-section of an Inconel 625 cylinder processed by SLS to 98.5% density. This specimen was post-processed by a HIP cycle consisting of 3 hours at 1240 ° C and 25000 psi. This cycle resulted in nearly full densification to 99.5%. A cross section of the HIPed specimen is shown in Figure 6 Specimen 051797, as SLS processed, 1.5% porosity. Temperature and pressure HIP maps (Ashby) were generated using software developed by one of the authors (Bourell).

Temperature and pressure HIP maps (Ashby) were generated using software developed by one of the authors (Bourell). The objective was to correlate experimentally observed density with predicted density. These maps show that for Inconel 625 parts with initial density in excess of 98%, the dominant mechanism of densification at temperatures higher than 1100° C and pressures in excess of 100 MPa (15 ksi) is power law creep. The maps predict that HIPing conditions of 1100° C and 100 MPa should be sufficient to densify Inconel 625 in excess86~' ..1-of 98% density in 1 hour.

Shown in Figure 8 is a cross-section of a PREP Ti-6Al-4V single layer (0.15 cm thickness) produced by SLS. This specimen successfully passed the leak test with a helium leak rate less than 1 X 10-10 standard cm3ts. Shown in Figure 9 are cross-sections of Ti-14AI-21Nb and Monel single layer specimens demonstrating full density capability.87I1300Figure 8 PREP Ti-6AI-4V, single layer, as SLS processedFigure 9 Single layer specimens of PREP Ti-14AI-21Nb (left) and Monel (right)CONCLUSIONSThe feasibility of SLSIHIP has been successfully demonstrated for a simple cylindricalshape made of Inconel 625 superalloy. Preliminary screening trials on PREP Ti-6Al-4V indicate that a gas impermeable skin can be produced by SLS. Research is ongoing to qualify additional materials including Monel, Ti-14Al-21Nb, Molybdenum and 17-4PH stainless steel. The next step in SLSIHIP process development will be to demonstrate a complex geometry part.

The next step in SLSIHIP process development will be to demonstrate a complex geometry part. SLSIHIP has tremendous potential for rapid, net shape manufacture of high performance metal components at reduced cost and shorter lead times.88ACKNOWLEDGEMENTSThe Laboratory for Freeform Fabrication at the University of Texas gratefully acknowledges funding support provided by the Defense Advanced Research Projects Agency and The Office of Naval Research and under DARPA/ONR contract NOOOI4-95-C-OI39 titled "Low Cost Metal Processing Using SLSIHIP". Lockheed Martin Vought Systems is the contract program manager.REFERENCESi Beaman, Joseph J. and Deckard, Carl R., Solid Freeform Fabrication and Selective Powder Sintering, 15th NAMRC, North American Manufacturing Research Conference Proceedings, 1987, pp. 636-640. 2 Deckard, C.R., Ph.D. Dissertation, Department of Mechanical Engineering, The University of Texas at Austin, 1988.

rect Selective Laser Sintering of High Temperature MaterialsG. Zong, Y. Wu, N. Tran, 1. Lee, D. L. Bourell, 1. J. Beaman, and H. L. MarcusCenter for Materials Science & Engineering and Department ofMechanical Engineering The University of Texas at Austin, Austin, TX 78712AbstractSelective Laser Sintering (SLS) involving a coexisting liquid and particulate solid during the SLS processing can be used to produce freeform parts directly with high temperature materials. Factors such as scanning laser power density, residence time, scan line spacing, the interfacial energies between the liquid and solid phases, powder bed biasing temperature, and sintering atmosphere greatly affect the microstructure evolution in the SLS process. Direct SLS of high melting temperature mixed powder materials was demonstrated using a 1.1 kW C02 laser SLS system. The relationship between the microstructures and the process parameters will be described.1.

The relationship between the microstructures and the process parameters will be described.1. IntrOductionSelective Laser Sintering is a pressureless sintering process used for Solid Freeform Fabrication (SFF). SLS has been used successfully to fabricate prototype parts and preforms with polymer, ceramic, or low melting temperature metallic materials [1-4]. A liquid phase is usually involved during the SLS processes. The atomic diffusion in the liquid phase is faster than that in the concurrent solid state process. The capillary force due to a wetting liquid is the driving force for rapid compact densification. The liquid also reduces the interparticle friction; thereby, aiding rearrangement of the solid particles. In addition, liquid dissolution of sharp particle edges and comers may allow more efficient packing. Thus, SLS processing control offers the opportunities for microstructure manipulations to optimize properties.

Thus, SLS processing control offers the opportunities for microstructure manipulations to optimize properties. With proper control of the above parameters, direct SLS of materials is possible. In this paper, direct SLS of high melting temperature mixed powder materials will be demonstrated using a 1.1 kW CO2 laser SLS system.2. ExperimentalTo perform selective laser sintering of high temperature materials, a high power laser selective sintering system was designed and assembled. There are four major components in the system: I) a 1.1 kw C02 laser system with focusing optics and scanning mechanism; 2) a72Sun station and its interface to control the laser pattern generation; 3) a sintering chamber with a resistance plate heater in part cylinder and inert gas purging capability; and 4) a powder delivery and leveling system.

Figure 1 shows schematic diagram of the sintering cell.The initial materials studied were Cu-Ni and WC-Co-Ni mixed powders as well as stainless steel 304L powders. Additive powders were also used. Optical metallograph, SEM, AES, EDS, and XRD were employed to study the sintering behaviors.x-v Movable 1.1kw CO 2 LaserLeveling DrumZnSeWindow"....Purging Gas Inlet...Gas OutletPowder FeedPart CylinderFigure 1 Sintering cell for high laser power SLS3. Results and DiscussionsThe main SLS processing concerns are the properties of the starting powder, the laser parameter settings, and the atmosphere in which powder sintering occurs. These parameters and their interactions determine the SLS product quality.

These parameters and their interactions determine the SLS product quality. The initial results of the SLS of high temperature materials using the system described above will be presented with evaluation of laser parameter effects, role of additives, and sintering atmosphere effects.3.1 Laser parameter effects in SLS73For a given SLS system, the laser parameters to be set in a typical SLS run are laser power density (P), scan speed (v), and scan spacing (S). Figure 2 shows the SLS sintering single layer thickness as a function of scan speeds and scan line spacing for 100 Ilm eu powder. The sintered layer thickness decreases with increasing scan speeds due to shorter interaction (sintering) time. This thickness also decreases with decreasing scan line spacing. Since the laser beam diameter of 4.5 mm is larger than the scan line spacings studied (0.254, 0.508, 1.016, 1.270 mm), scan overlapping occurred in these tests. The smaller the scan line spacing, the more scan overlapping.

The smaller the scan line spacing, the more scan overlapping. Therefore, the sintered layer thickness decreases with increasing scan overlapping. This relationship can be explained by considering the fact that the thermal conductivity and reflectivity of the sintered solid are higher than those of the starting powder. The more the scan overlapping in a laser scan, the more laser energy will be transferred away by heat conducting through the sintered solid and reflected away by sintered solid surface. Thus the sintering layer thickness decreases.6. 8=0.254 mm 08=1.27 mm08=0.508 mm08=1.016 mm.9E .s.8(J) (J) a> c:: .::s:. .2 .c:: l- I - a> ~ ....I.7.6.5jJ.4.3246812 8can 8peed (mm/sec)1014161820.Figure 2 SLS sintering layer thickness as a function of laser scan speeds and scan line spacing.

Laser power = 2oow, beam size=4.5 mm, nitrogen flow rate = 60 1/min, no powder bed biasing temperature, for copper powder (-1OOllm) material.To obtain a desired microstructure, sintering laser power (temperature) and scan speed (time) are the most important factors. The quality of the blend and analysis of the atmosphere also affect the microstructure. SLS tests and metallographic studies have been made to investigate the effect of laser power and scan speed on the sintering of a Bronze+Ni system. Figures 3 shows the typical sections of the SLS sintered part. These pictures show the strong effect of laser power and scan speed on the amount, size, and shape of residual porosity. The7422sintered porosity is lower at higher laser power and lower scan speed. The pores become smaller and more rounded as the laser power increases and scan speed decreases because both sintering temperature and time are increased, more rearrangement [5] is involved during sintering.

The meta1lographic study confirms the effect of laser power and scan speed. Figure 4 shows undersintered microstructure at conditions of Figure 3(a) and well sintered microstructure at conditions of Figure 3(c). Successful parts (Figure 5) with -82% of theoretical density were obtained at laser power = 280w, beam size=4.5 mm, scan speed = 4.2mm1sec, scan line spacing =0.5mm, layer thickness=0.25mm, purging nitrogen flow rate = 60 I fmin, powder bed biasing temperature=350 °C, for 60 wt% Bronze 9010 (-I00Jlm)-30 wt% Ni (-100Jlm)-Cu3P (63Jlm) powder mixture.During SLS process, laser heating causes temperature gradients that lead to the formation of residual stresses and strains throughout the sintered part. In the extreme case, these residual stresses will cause sintered layer curling. Laser power density and scan speed, together with scan path, part size, and powder bed biasing temperature, have strong effects on the generation of residual stresses in the sintered parts.

Laser power density and scan speed, together with scan path, part size, and powder bed biasing temperature, have strong effects on the generation of residual stresses in the sintered parts. Initial results showed that curling can be overcome or reduced by suitable setting these parameters. These results will not be discussed in detail in this paper.3.2 Role of CU3P in SLS of Cu-Ni materialsTests have been carried out with several kinds of Cu-Ni and Bronze-Ni powders containing various amounts OfCu3P (0-10 wt%). These tests show that the present success of SLS ofCu Ni powders is, to a large extent, dependent upon the addition of CU3P. This section will discuss the role of Cu3P in SLS of Cu-Ni materials.X-ray analysis of the starting powder mixture of 60 wt% Bronze-30 wt% Ni -CU3P and its SLS sintered part indicated that CU3P peaks showed up on the profile of the starting powder but not on the profile of SLS sintered samples (Figure 6).

However, EDS analysis (Figure 7) showed that P is still in the sintered part. This phenomena can be explained by referring to the Cu-P equilibrium diagram [6] in Figure 8. The Cu-CU3P partial system is a simple eutectic. The solubility of P in Cu is about 1.7 wt.% at the eutectic temperature of 7140C. At the usual SLS sintering temperatures (above the melting point ofcopper, 1085 OC), CU3P (melting point 10220C) will melt and there is considerable solubility ofP in the molten copper. When CU3P melted, P is diluted by the copper in bronze. Calculation indicates that total weight percentage of P in copper becomes 1.27%, about the solubility of P in copper at room temperature. Thus, after sintering, P is a solute in the Cu-P solution instead of forming CU3P.

Smooth sintering layers (less balling) can be obtained when CU3P is added to the powder mixture. Therefore, the present of P in the SLS process may enhance sintering by improving the wettability of Cu-Ni system through reducing surface tension of liquid copper. Since the free energy of formation of P oxide is much smaller (more negative) than that of Cu and Ni, P may also act as a localized oxygen 'getter' and partially prevent the oxidations of Cu and Ni, thus enhance sintering. In-situ Auger fracture analysis of the fracture surface of SLS sintered 60 wt% Bronze-30 wt% Ni -Cu3P sample revealed strong oxidation and high P content in the oxide (Figure 10), though no wetting problems occurred during the sintering process. The carbon peak originates from the carbon on the free surface of the pores. This analysis supports the idea that P is a localized oxygen 'getter'.

The carbon peak originates from the carbon on the free surface of the pores. This analysis supports the idea that P is a localized oxygen 'getter'. Systematic investigations into the effects of P and other wetting additions on the sintering behavior of Cu-Ni system in an oxygen clean environment are highly desirable.3.3 Sintering atmosphere effects in SLSThe sintering atmosphere is one significant factor affecting selective laser sintering, especially high temperature SLS. The atmosphere protects against surface contamination during sintering. Additionally, a cleaning function is desired from the atmosphere to remove films (for example oxides) for rapid melt flow. The atmosphere plays an important role in reducing surface films which can alter the wetting angles. In conventional liquid phase sintering, best densification and properties can be achieved using vacuum sintering in many systems [5]. Residual atmosphere can be trapped in pores which seal during liquid flow.

Residual atmosphere can be trapped in pores which seal during liquid flow. Such trapped atmosphere inhibits full densification. Sintering in an inert or insoluble atmosphere is most detrimental for this reason. The present SLS system uses an inert gas purging system to reduce the surface layer contamination. This section will address some problems associated with such an approach.Residual gas analysis (RGA) was performed to analyze the gas in the working chamber. With 60 I fmin nitrogen purging, which is the setting for all tests discussed in the paper, the N2 to 02 partial pressure ratio 5000: 1 was measured. SLS of stainless steel 304L powder was unsuccessful using such an environment due to the oxidation problem.SLS of WC-Co-Ni system produced strong parts with rough surfaces (Figure 11). Wetting angle of WC by both liquid Co in hydrogen and liquid Ni in vacuum at 15000C is 0 [7]. However, the oxidation of Co and Ni causes poor wetting and thus balling.

Wetting angle of WC by both liquid Co in hydrogen and liquid Ni in vacuum at 15000C is 0 [7]. However, the oxidation of Co and Ni causes poor wetting and thus balling. Auger surface analysis of sintered 70 wt.% (WC-12Co)-20 wt.% Co- Ni showed oxygen in the sample, as seen in Figure 12. For a sintered layer, balling occurs on the top of the layer, and the bottom of the layer shows good wetting. Optical studies of the bottom of the layers revealed expected composite microstructure (Figure 13a), and WC wets well by metal matrix Co and Ni. The81study implies that the oxidation of Co and Ni occurs only on the liquid surface during the SLS process. Figure 13b shows the fracture surface of the sintered WC-Co-Ni sample, where voids with liquid pools are obvious. These voids may be introduced by N2 gas trapped in liquid during liquid flow in SLS processing.Figure 10 AES analysis of the in-situ fracture surface of the SLS sintered 60 wt% Bronze-30 wt% Ni -CU3P.4.

SummaryDirect Selective Laser Sintering of high temperature powder mixtures has been demonstrated. Initial results indicate that SLS involving a coexisting liquid and particulate solid during the sintering can be used to produce freeform parts directly with high temperature materials. Many factors including scanning laser power density, residence time, scan line spacing, powder bed bias temperature, atmosphere, interfacial energies between the liquid and solid phases, and roles of additives are discussed in terms of the microstructure evolution.821 I iThese parameters and other factors such as solid phase solubility in the liquid, relative penetration of the liquid, particle size and distribution, powder leveling thickness and density are still to be studied.Figure 11 SLS fabricated WC-Co-Ni parts.wElectron Energy eVFigure 12 AES analysis of the (WC-12 wt.% Co)-20 wt.% Co-IO wt.% Ni sample, sputtering time = 6min.83Figure 13 wt. % Co-lO wt.

% Co-lO wt. % Ni sample.(a) Optical and (b) SEM micrographs of SLS sintered (WC-l2 wt.% Co)-20AcknowledgementThis research was supported by DTM Corp. at Austin.84References1. 1990 Solid Freeform Fabris;ation Symposium PrOkeeginaS, Edited by J.J. Beaman, H.L. Marcus, D.L. Bourell, J.W. Barlow, 1990, The University of Texas at Austin, Austin, Texas.2. 1991 SQlig FrpefQrm Fabrication Symt1Qsium PrOkeedinaS, Edited by H.L. Marcus, J.1. Beaman, J.W. Barlow, D.L. Bourell, R.H. Crawford, 1991, The University of Texas at Austin, Austin, Texas.3. H. L. Marcus, J. J. Beaman, J. W. Barlow, and D. L. Bourell, "Solid Freeform Fabrication: Powder Processing," Ceramic Bulletin, 69(6), 1990, 1030..31.4. D. L. Bourell, H. L. Marcus, J. W. Barlow, and J. J. Beaman, "Selective Laser Sintering of Metals and Ceramics," to be published in the October 1992 issue of the International Journal ofPowder Metallurgy.5. Randall M.

Marcus, J. W. Barlow, and J. J. Beaman, "Selective Laser Sintering of Metals and Ceramics," to be published in the October 1992 issue of the International Journal ofPowder Metallurgy.5. Randall M. German, LiQUig Phasp Sintprina (New York, NY: Plenum Press, 1985), 127.. 151.6. "Binary Alloy Phase Diagrams," second edition, ASM International, 1990 Vol. 2, T. B. Massalski, editor..in..chief.7. G. S. Upadhyaya, "Wetting of Ceramics by Melts .. An Electronic Approach," Sintprpg Mptal"CeramiS; CQmposites, edited by G. S.. Upadhyaya, Elsevier Science Publishers B.V., Amsterdam 1984, p. 41.85.

RECT SELECTIVE LASER SINTERING OF REACTION BONDED SILICON CARBIDESebastian Meyers*, Jean-Pierre Kruth*, Jef Vleugels **KU Leuven, Department of Mechanical Engineering, division PMA, Celestijnenlaan 300, 3001 Heverlee, Belgium ** KU Leuven, Department of Materials Engineering, Kasteelpark Arenberg 44, 3001 Heverlee, Belgium AbstractThree-dimensional reaction bonded silicon carbide (SiSiC or RBSC) parts have been produced by direct selective laser sintering (SLS). Unlike previously investigated processing routes, which make use of a sacrificial polymer binder to form green parts, the parts in this work are built by scanning subsequent layers composed of a mixture of silicon and silicon carbide powders. A fibre laser is used to selectively melt the silicon under an inert argon atmosphere, resulting in porous preforms of sufficient strength for further handling and processing.

A fibre laser is used to selectively melt the silicon under an inert argon atmosphere, resulting in porous preforms of sufficient strength for further handling and processing. After impregnation with a graphite suspension and infiltration with liquid Si at 1450°C, highly dense reaction bonded silicon carbide parts are obtained.IntroductionSilicon carbide is an engineering ceramic which is difficult to manufacture by most conventional powder processing routes [1]. Reaction bonding, however, is a process that allows for near net- shape production of SiC materials at lower processing temperatures and in shorter times [2]. It relies on silicon infiltration of a carbon-containing preform and the subsequent reaction to SiC. The final product is reaction bonded silicon carbide (RBSC), i.e., a matrix composite containing primary SiC, secondary (reaction formed) SiC and residual Si.

The final product is reaction bonded silicon carbide (RBSC), i.e., a matrix composite containing primary SiC, secondary (reaction formed) SiC and residual Si. This residual Si is the main drawback of the processing route, since it limits the maximum service temperature of the material (Si melts at 1420°C). Nevertheless, RBSC has some interesting properties like high hardness, good thermal conductivity, chemical resistivity and low thermal expansion, which make it suitable for industrial applications.In order to facilitate the production of RBSC for industrial applications, additive manufacturing techniques can be used. Selective laser sintering (SLS) and selective laser melting (SLM), for example, are powder bed fusion layered manufacturing techniques which can produce complex 3D parts [3]. The selective laser sintering of RBSC using a sacrificial polymer binder has already been extensively studied. Vail et al.

The selective laser sintering of RBSC using a sacrificial polymer binder has already been extensively studied. Vail et al. [4] produced polymer encapsulated SiC particles and fabricated green parts with good strength and relative densities of 48 – 51%. Further research by Evans et al. [5] [6] focussed on using a char-yielding binder in order to provide residual carbon in the structure. The residual carbon could then react with molten Si during infiltration to form a secondary SiC phase. This approach resulted in highly dense RBSC parts. However, Stevinson et al. [7] observed that the surfaces showed liquid Si over-extrusions after infiltration. Hon et al. [8] performed selective laser sintering of SiC-polyamide powder mixtures. The effect of the SLS parameters, energy density and the initial powder blend composition on the produced parts were studied.Some authors have also attempted to produce RBSC parts without using a sacrificial polymer binder. Birmingham et al.

Birmingham et al. [9] showed that selective laser (reaction) sintering of Si powder beds in a C2H2 gas atmosphere could result in single- and multilayer shapes with high SiC content. Streek et al. [10] used laser micro sintering to process both SiSiC with a q-switched laser and pure SiC powders with a continuous (cw) laser. Finally, Vacuher et al. [11] did not focus on RBSC, but produced aluminium – SiC metal matrix composites by selective laser sintering blended powder mixtures with an Nd:YAG q-switched laser. This resulted in final parts with porosities of 40- 60%.In the present work, porous preforms were produced by direct selective laser sintering blended Si-SiC powder mixtures without using a sacrificial polymer binder. Partial melting of the silicon acted as the powder consolidation mechanism. After SLS, the preforms were impregnated with a graphite suspension to provide a carbon source during subsequent liquid Si infiltration.

After SLS, the preforms were impregnated with a graphite suspension to provide a carbon source during subsequent liquid Si infiltration. This allows for in situ formation of secondary (reaction formed) SiC. After Si infiltration, highly dense SiSiC parts are obtained. A schematic of the process is shown in Figure 1. Density measurements and microscopy imaging are used to examine the final quality of the products.Powder PreparationSelective Laser SinteringCarbon ImpregnationReactive Silicon InfiltrationFigure 1: processing schematic: direct SLS of SiSiC.Experimental ProceduresPowder PreparationPowder mixtures were prepared by mixing 67 weight% silicon carbide (SiC, CARBOREX BW F320, Washington Mills, purity = 99.2%, d50 = 29 µm) and 33 weight% silicon (Si, SIMET 985, Keyvest, purity = 98%, d50 = 45 µm) in a polyethylene container on a Turbula mixer for 6 hours at 75 rpm.

The mixing was done dry and resulted in a homogeneous SiSiC powder that could subsequently be loaded into the SLS machine. The starting powders are shown in Figure 2.Figure 2: Washington Mills SiC powder (CARBOREX BW F320, left) and Keyvest Si powder (SIMET 985, right).Selective Laser SinteringSelective laser sintering was done on a commercial Concept Laser Mlab cusing machine, equipped with a 100 W continuous (cw) fibre laser with a spot size of 50 µm. The parts were built in various batches of several 10x10x10 mm³ cubes. The scanning parameters were varied across these cubes in order to obtain an optimal parameter set. Laser powers were varied between 10 – 95 W and scanning speeds from 30 – 800 mm/s. The scanning was done continuously with a 90 degree rotation between layers. The powder layer thickness was fixed at 60 µm and no preheating was used.Carbon ImpregnationIn order to impregnate the SLS preforms with carbon, a graphite-acetone suspension was used.

The powder layer thickness was fixed at 60 µm and no preheating was used.Carbon ImpregnationIn order to impregnate the SLS preforms with carbon, a graphite-acetone suspension was used. To this end, 20 vol% of graphite powder (C, Dragon Seal Graphite MF2/99.5-99.9RG, NGS Naturgraphit GmbH, purity = 99.5-99.9%, d50 = 2µm) was suspended in acetone (C3H6O, Nyssens Graphics, purity = 99%) and mixed using a magnetic stirring device for 30 minutes. After this, impregnation was accomplished by dipping the preforms in the suspension twice for 5 minutes with an intermediate drying step.Reactive Silicon InfiltrationAfter graphite impregnation, the preforms were transferred to a graphite crucible. The crucible was lined with graphite paper on the inside and contained silicon wafer chunks (Imec, Leuven) on the bottom. The preforms were placed on top of the silicon wafer chunks.

The crucible was lined with graphite paper on the inside and contained silicon wafer chunks (Imec, Leuven) on the bottom. The preforms were placed on top of the silicon wafer chunks. The crucible containing the preforms was transferred to a hot press (FCT Systeme , model W100/150-2200-50 LAX) for high temperature silicon infiltration under vacuum (10-1 mbar). The temperature was increased to 1450°C at a heating rate of 50°C/min. A dwell time of 30 minutes was applied at 1450°C and the crucible was subsequently cooled down to room temperature.MeasurementsAfter fabrication, different measurements were done to evaluate the quality of the cubes. Density measurements were done on the porous preforms and the final RBSC composites using the Archimedes method (Acculab atilon ATL-244-1). X-ray diffraction measurements (Seifert 3003 TT) were carried out to gain information on the phase composition of the preforms.

X-ray diffraction measurements (Seifert 3003 TT) were carried out to gain information on the phase composition of the preforms. Scanning Electron Microscopy (SEM, FEI XL30-FEG) was done on the top and side surfaces of the porous preforms, and finally light microscopy (Leica DMILM HC) was performed on the final RBSC composites to assess the microstructure and SiC content.Results & DiscussionSelective laser sintering was done with different parameter sets. High power – high scanning speed sets were used at the start, but resulted in fragile and dimensionally unstable preforms. More robust and dimensionally accurate preforms were obtained by using parameter sets with reduced laser power and scan speed.

More robust and dimensionally accurate preforms were obtained by using parameter sets with reduced laser power and scan speed. Parameter sets can be compared by expressing the laser energy density:𝐸 =𝑃 𝑣 ∗ 𝑠 ∗ 𝑙With:E = laser energy density in J/mm³ - P = laser power in W - v = scanning speed in mm/s - - s = scan spacing or hatch spacing in mm l = powder layer thickness in mmPreforms built with similar laser energy densities, but with a higher P and v, were found to be of lower quality than preforms with a lower P and v, as illustrated in Figure 3. The dimensions of the preforms were as-defined.Figure 3: High power - high scanning speed (P = 40W, v = 200 mm/s, left) versus low power - low scanning speed (P = 20W, v = 100 mm/s, right) with comparable E ≈ 43 J/mm³.The effect of scan spacing was investigated by investigating top surfaces of different preforms with scanning electron microscopy.

A comparison of two preforms with different scan spacings is presented in Figure 4. The preform on the left was made using a scan spacing 2 times higher than the preform on the right. The surface shows some protrusions, which can be linked to the scan tracks (arrows, with 90 degree rotation). The preform with smaller scan spacing does not show these features, but clearly still contains some porosity.Figure 4: SEM images of SLS preform top surfaces with P = 20W, v = 90 mm/s (a: s = 77 µm, b: s = 35 µm).In order to quantify the porosity, the density was measured according to the Archimedes method. These measurements showed that the densities vary between 40 and 50% of the theoretical density and decrease when the laser energy density increases. This leads to the conclusion that low laser energy densities (30-100 J/mm³) in combination with low laser powers and scanning speeds lead to better preforms.

This leads to the conclusion that low laser energy densities (30-100 J/mm³) in combination with low laser powers and scanning speeds lead to better preforms. These preforms, although porous, are dimensionally accurate and strong enough to be processed further.X-ray diffraction (XRD) was performed in order to identify the phases present in the fabricated preforms. The measurements showed that the preforms consist of silicon carbide (α-SiC, polytype 6H) and silicon. A typical XRD pattern is shown in Figure 5.35003000SiC 6HSi2500200015001000500020304050607080902θ (deg)Figure 5: XRD spectrum of a Si/SiC composite after SLS.After impregnation with a graphite suspension and infiltration with liquid silicon at 1450°C, RBSC cubes were obtained. These cubes showed some surface protrusions of infiltrated silicon (over-infiltration). However, they appeared to be highly dense. The final microstructure is shown in Figure 6.

These cubes showed some surface protrusions of infiltrated silicon (over-infiltration). However, they appeared to be highly dense. The final microstructure is shown in Figure 6. The part was scanned with parameters comparable to the one from Figure 4.a, albeit with a slightly lower scanning speed (P = 20W, v = 70 mm/s, s = 77 µm). The bright matrix phase is silicon, whereas the grey contrast phase is α-SiC starting powder. Around the α-SiC, a dark grey border of reaction formed SiC can be seen. This is shown in higher magnification on Figure 7. The reaction formed SiC is a result of the reaction between the liquid Si and the impregnated graphite (C) during infiltration:Si (l) + C (s)  SiC (s)Figure 6: Typical microstructure of a RBSC part produced by SLS and subsequent post-processing.Figure 7: Detailed view of the dark grey reaction formed SiC around the primary SiC grains.Clearly, the final RBSC cubes still contain a high amount of silicon.

This is due to a too low amount of carbon in the preform after graphite impregnation. This issue can be addressed in one of the following ways:Carbon impregnation under external pressure should yield a higher carbon content in thefabricated preforms; Impregnation with a phenolic resin and subsequent de-binding could result in an increased carbon content;Carbon could be added to the base powder mixture so that the preforms after SLS contain a certain amount of carbon, or better yet a higher amount of silicon carbide thanks to the reaction of molten silicon with free carbon.In order to quantify the current amount of SiC, image analysis software was used on 10 different light microscopy images. This showed that the cubes scanned with a high scan spacing of 77 µm contain only 35-40% SiC. From this, a theoretical density of 2.66 g/cm³ can be calculated.

This showed that the cubes scanned with a high scan spacing of 77 µm contain only 35-40% SiC. From this, a theoretical density of 2.66 g/cm³ can be calculated. Cubes scanned with a lower scan spacing of 50 µm contain more SiC (60%), resulting in a theoretical density of 2.85 g/cm³.The densities of the infiltrated RBSC shapes were compared to the calculated theoretical density and found to depend on the used laser scan spacing. A larger scan spacing resulted in relative densities of 95%, whereas a lower scan spacing yields significantly lower final densities of 70%. These lower densities are possibly caused by a large volume of closed porosity in the middle of the part, which effectively inhibits liquid silicon infiltration. A cross section of a low scan spacing part is shown in Figure 8. The lighter grey parts are infiltrated with molten Si. Clearly, a core-shell like structure was obtained.

A cross section of a low scan spacing part is shown in Figure 8. The lighter grey parts are infiltrated with molten Si. Clearly, a core-shell like structure was obtained. The silicon could successfully infiltrate the edges of the part, but did not enter the core.Infiltrated areaNon infiltrated coreFigure 8: Cross section of a Si-infiltrated low scan spacing SLS part with shell-like structure.The dimensions of the final infiltrated Si-SiC cubes were, again, as-defined. This means no shrinkage occurred during post processing, which was to be expected since there was no polymer de-binding step involved.Conclusions & SuggestionsRBSC parts have been produced without the use of a sacrificial binder by direct selective laser sintering. During direct SLS, melting and re-solidification of the silicon could act as a binder mechanism. High laser power – high scanning speed SLS parameter sets failed to produce successful parts.

During direct SLS, melting and re-solidification of the silicon could act as a binder mechanism. High laser power – high scanning speed SLS parameter sets failed to produce successful parts. However, low laser power – low scanning speed parameters resulted in Si-SiC preforms of sufficient strength for further handling. After carbon impregnation and liquid silicon infiltration, highly dense (~95%) RBSC parts were obtained. These parts contained a high amount (~60%) of residual silicon, but strategies to resolve this issue were proposed and are currently under investigation.AcknowledgementsThis work was financially supported by the Research Foundation – Flanders (FWO) under project FWO-G.0956.14N. The authors acknowledge support of the research fund of KU Leuven through Project GOA/15/012-SUMMA and wish to thank Ms. Raya Mertens and Mr. Xia Yu for their assistance.References[1] K. Biswas, „Solid State Sintering of SiC-Ceramics,” Materials Science Forum, vol.

rect Selective Laser Sintering of SuperalloysFritz Klocke and Christian Wagner Fraunhofer-Institute of Production Technology IPT Aachen, GermanyAbstractThe advantages of powder metallurgy lie within the large degree of freedom for material design and thus is especially used in the production of high performance parts. Layer manufacturing is an appropriate method to produce complex parts rapidly. Direct Selective Laser Sintering (SLS) presents a technology which combines both benefits. Therefore many efforts are done today to qualify new materials for SLS [DAS 98, MEI 99, OVE 01, WOL 00]. Particulary materials, which are hard to cut, to cast or to shape in any else matter, are of interest.In the presented paper investigation results on Selective Laser Sintering of metals are shown on the basis of the nickel base alloy INCONEL™ 718. First, a process model has been created to discribe the mechanisms of SLS of metals.

First, a process model has been created to discribe the mechanisms of SLS of metals. On the base of the model, process simulations and experimental investigations have been performed. In some cases, affiliating a heat treatment after the laser sintering step is favorably to improve the metallic structure and thus has also been tested. Finally, the metallograhic structures and mechanical properties were analysed.1871. IntroductionIn Selective Laser Sintering of high performance metals, two bad effects have been observed: balling and oxydation. During process observation, it was found that materials without a semisolid phase, like pure metals, show strong balling effects, whereas alloys with a semi-solid phase are better to process. Therefore it was assumed, that metal droplets in a superliquid state tend to flow together abruptly and in a semisolid phase the fusion process lasts longer, which is benefical for the SLS process.

Therefore it was assumed, that metal droplets in a superliquid state tend to flow together abruptly and in a semisolid phase the fusion process lasts longer, which is benefical for the SLS process. If this theses would apply, the most important job during SLS would be to heat the process zone exactly between the liquidus and solidus temperature, when processing technical alloys.To verify the theses, first a base test with two metal grains was done. Then a process model, which enables the prediction process parameters, was developed. The process model was the base for an FEM- simulation of SLS. It also gives hints for process boundary conditions, like maximum oxygen contents in process atmosphere. The results and conclusions were used to build sample parts of IN 718.2. Base test: sintering behaviour of 2 grainsThe objective of this test was to determine the melting behaviour and fusion velocity of semisolid and non-semisolid metals.

Base test: sintering behaviour of 2 grainsThe objective of this test was to determine the melting behaviour and fusion velocity of semisolid and non-semisolid metals. The test constallation of the 2-grain test is shown in fig. 1a). Two grains were positioned side by side on a polished aluminium plate. A laser beam with controlled power P, which has been varied during testing from 125 W to 20W, and. was occuring vertically during a exposure time texp. = 1000 ms on the particles. The objective was to observe how long the fusion time tsint. lasts for different laser powers. For the description of the sintering stage, the coalescence ratio S was introduced, see fig. 1b).a)b)laser beamS = 0 %a0coalescence ratio S:S=a0−a0( ta Sint ∞− a)%100powder particlealuminium platevolume constance:0 % < S < 100 %a =∞1 23 2a0a(tSint.)Test paramenters – CO2-laser power P: 20 - 125 W var.

– laser beam diameter: 0,5 mm – laser beaming time texp.: 1000 ms Test materials – stainless steel (1.4404), copper (Cu99) – particle size 100-200 µmS = 100 %a∞Fig. 1:a) test set-up and test programme; b) explanation of introduced test dimensions188The process was recorded with a high speed camera with a picture frequency of 2000 Hz. Tested materials have been copper of 99 % purity (Cu99) and stainless steel 1.4404. Copper has its melting point at Tmelt = 1083 °C, whereas the stainless steel starts to melt at Tsol. = 1375 °C and is completely molten at Tliq. = 1430 °C, see table 1.materialcomposition [%]TsolidTliquid∆Tliquid-solidcopper Cu99Cu 99div.

The coalescence behaviour of stainless steel is different: the sintering time of stainless steel 1.4404 depends on the laser power P, see fig. 3. The coalescence ratio S has been determined by evaluating the high speed photographs from the tests. For laser powers of 50 W and up, that the fusion times tsint. to complete coalescence (S = 100 %) are all very short (tsint.< 5ms). In fig. 2b), the coalescence process of stainless steel for a laser power of 30 W is shown. The final stage is reached within approx. 40 ms. With decreasing laser power P, the coalescence time tsint. grows. The fusion times for 1.4404 range from 2 ms to aprox. 70 ms.Referring to these observations, the theses of the sintering in the semisolid stage is applying: in the semisolid state, the viscosity is higher than in the superliquid state and the fusion duration is longer. In fig. 3, the curves of copper are not included.

In fig. 3, the curves of copper are not included. They have all the same formation and can be compared to the stainless steel graph for P = 125 W. The laser power of 40 W, which has been applied in the test in fig. 2a), is also the lowest power, where sintering of copper occurs.IN 718 has not been included in the 2 grain test, but it is assumed that the results can be transferred from 1.4404 to other semisolid alloys. A justification for this assumption is seen in the fact, that both tested materials, Cu99 and 1.4404, flow together within the same time (≈ 2ms), when being superliquid. On the other hand, the superliquid viscosity µ of a Fe-Cr20-Ni10,5 alloy amount to 9,2 mPa⋅s at 1500 °C, whereas Cu has viscosity of 3 mPa⋅s at 1100 °C [KAW 88]. As a result, a difference in viscosity within this order of magnitude has no big impact. The viscosity in the semisolid range, which is considered as a decisive factor, must be tremendously higher than in the superliquid area.189Fig.

The viscosity in the semisolid range, which is considered as a decisive factor, must be tremendously higher than in the superliquid area.189Fig. 2:a)tsint. = 0 msb)tsint. = 0 mstsint. = 0,5 mstsint. = 10 mstsint. = 1,0 mstsint. = 20 mstsint. = 1,5 mstsint. = 30 mstsint. = 2,0 mstsint. = 40 msHigh speed recordings of the 2-grain tests: a) Copper Cu99, P = 40 W; b) stainless steel 1.4404, P = 30 W190100 10080 80125 Wlaser power P]%50 W[ S o i t a r60 60e c n e c s e l a o c40 4040 W30 W20 W20 200 00 010 1020 2030 3040 4050 5060 6070 70sintering time tsint. [ms]Fig. 3:Coalescence times for 2 stainless steel particles in dependance of the laser power P3. Modeling of SLSWith respect to the 2-grain test results, the temperature field during SLS should be kept between Tsol. and Tliq. within the actual powder layer and slightly below.

Modeling of SLSWith respect to the 2-grain test results, the temperature field during SLS should be kept between Tsol. and Tliq. within the actual powder layer and slightly below. For that purpose, a process model has been created to calculate the temperature distribution within the process zone. The base of this model is a standard material cell which represents the typical material behaviour of the powder bed during SLS. The material properties and physical pocedures during laser sintering can be asorted to 4 different categories: the thermal sub-process, the metallurgical sub-process, the chemical sub-process and the sintering sub-process. Each of these 4 processes was treated separately at first. After that the submodels had been composed to an integrated process model.3.1 The thermal submodel The thermal submodel includes all thermal dimensions of the process, e.g. the specific heat cp and the thermal conductivity λ. Both properties are temperature dependent.

the specific heat cp and the thermal conductivity λ. Both properties are temperature dependent. Tabled values of cp and λ for a large range of technical alloys can be found in [VDI 97]. The specific heat is only a function of the temperature, whereas the thermal conductivity is a function of the temperature and the contact interface between the particles. The latter could also be expressed by the relative density ρrel. as will be demonstrated later. The basic equation for the heat transfer problem is then:191( ρρ rel)( ) Tc p∂ T ∂ t=( ρλ , T rel)∂ ∂2 T (cid:1) 2 neq. 1For powder beds during laser sintering, ρ is a function of the void fraction respectively of the relative density ρrel.. Due to the dependancies in eq. 1, the heat transfer problem becomes non-linear and can only be solved by numerical methods.The thermal submodel also includes the definition of the heat source term, which has to be added to eq. 1.

1. Two important factors in defining the heat source are the laser energy absorption and optical penetration depth. The absorption grade A has been measured by the method of diffuse reflection (DR) with an FT-IR spectrometer 1700X from Perkin-Elmer Ltd.. It was determined to 18-20% for IN 718. For the calculations, the value was set to A = 0,2. The optical penetration depth δopt. for metals is rather low with 42 nm < δopt. < 106 nm for a temperature range of 300 K < T < 1800 K [HOF 93]. Therefore a surface heat source can be considered when calculating laser assisted heat-up procedures of metals. To take into account the penetration in the powder bed through the particle gaps, a penetration factor of δ´opt. = 0,15 µm, which is half of the average diameter of the powder particles, was included.

= 0,15 µm, which is half of the average diameter of the powder particles, was included. The equation for a laser beam with gaussian energy distribution as volumetric heat source becomes then:( zyxq ,,)−=A δ ´ opt.I0exp   −2x2+2 r by2−z δ ´ opt.   ,eq. 2where I0 is the intensity in the beam center and rb the beam radius.Also the latent heat Hmelt. of the crystalline metal material is considered in the calculations. The exact value for IN 718 has been unknown, so that the latent heat for nickel was taken Hmelt.(Ni) = 289 J/g. In the calculations, a linear assignment of the melting energie between Tsol. and Tliq. was assumed.3.2 The metallurgical submodel This submodel considers the metallurgical material properties of the metal, e.g. the solidus and liquidus temperature Tsol. and Tliq. or the values of the latent heat of melting or of lattice transformation.

the solidus and liquidus temperature Tsol. and Tliq. or the values of the latent heat of melting or of lattice transformation. The metallurgical submodel is also important for explanations of the processes in the cooling period and the resulting metallic structure. The metallurgical process might become even more relevant, if the process of melting and the refering viscosity in the semisolid state would be investigated in more detail. When studying laser sintering of powder mixtures, the process of microalloying also requires metallurgical modelling by means of phase diagramms.3.3 The chemical submodel The chemical submodel describes every process of chemical material transformation. The most important and worst reaction for metal laser sintering is oxydation. Oxyde skins prevent proper sintering and wetting [SCH 92]. The general reaction equation for a metal oxydation is:192Me+1 2{ O 2}→MeO, with(cid:1)∆G < 0 ,eq. 3where ∆G is the free enthalphy.

The general reaction equation for a metal oxydation is:192Me+1 2{ O 2}→MeO, with(cid:1)∆G < 0 ,eq. 3where ∆G is the free enthalphy. If ∆G is ≥ 0, the oxydation is impossible, if ∆G is < 0, the oxydation could happen. The bigger the absolute value of ∆G is, the more likely the reaction is. The free enthalpy can be calculated from∆=∆GG0+RTln1 ⋅ 2/1 p O2aMe−=RTlnKMeO+RTln1 ⋅ 2/1 p O2aMe<0eq. 4Values of the free creation enthalphy ∆G0 and of the equilibrum constant K for a large variety of elements and bondings can be found in [BAR 73]. When an element is present in a diluted form, like in alloys, the chemical reactivity, expressed by the activity a, changes. For alloys containing many elements, like for example for IN 718, the acticity aMe of each element can be set to 1. The graphs for metals and the referring oxygen partial pressures can be derived from an Richardson-Ellingham diagram [FRO 94].

The graphs for metals and the referring oxygen partial pressures can be derived from an Richardson-Ellingham diagram [FRO 94]. For IN 718 the most critical element is aluminium. Following the Richardson-Ellingham diagram, the partial pressure of oxygen pO2 must be less than 10-25 atm to prevent oxydation.3.4 The sintering submodel The sintering submodel includes mainly rheological processes. One aim is to determine the differential equation of the fusion process of two plastified particles, depending on the surface tension and viscosity. Another objective is the transfer of the base case of two particles to a package of more particles. The purpose of the sintering submodel is also to explain clustering of particles during sintering (balling effect).The inital powder bed density has been determind to ρrel.(IN 718) = 0,55. Following the 2-grain tests, the powder bed becomes fully dense within 2 ms in zones where the temperature exceeds Tliq..

Following the 2-grain tests, the powder bed becomes fully dense within 2 ms in zones where the temperature exceeds Tliq.. The sintering stage and the referring relative density and contact area is dependent on temperature and time. To mount an equation of the kind ρrel. (T, t), more detailed investigations of isothermal sintering rates would be necessary. For the calculation carried out here, a linear raise of the relative density from 0,55 to 1 in the heat-up period from Tliq. to Tsol. has been assumed.For the thermal conductivity, the size of the contact interface of the particles is relevant. The contact between the powder particles is setting the conductive cross section of the material continuum. It ranges from a point contact of the spheric particles to a 100% contact, see fig. 4. In measurements, the thermal conductivity λ of the powder bed has been determined after the laser flash analysis (LFA) method using a LFA 427 system from the Netzsch GmbH.

4. In measurements, the thermal conductivity λ of the powder bed has been determined after the laser flash analysis (LFA) method using a LFA 427 system from the Netzsch GmbH. At room temperature, it was found that λpowder (IN 718)(20 °C) is less than 1 W/mK. The value for compact IN 718 material is λ(IN 718)(20 °C) = 13 W/mK and for Tsol. approx. λ(IN 718)(1200 °C) = 33 W/mK [VDI 97]. Thus, λ has been set to:193T < Tsol.:1=λW mKTsol. < T < Tliq.:=λ1(+− TT sol − T T.liq.sol.)32W mKT > Tliq.:33=λW mK3.5 The integrated model The above mentioned subprocesses have been consolidated to a complete model. For this consolidation, a standard material cell has been created, see fig. 4. The standard material cell was developed from a powder particle, lying in a cubic powder package. The process stages are heating-up, melting start and wetting the underlying particle, coalescence and the final stage.

The process stages are heating-up, melting start and wetting the underlying particle, coalescence and the final stage. In the wetting stage, the underlying particle is heated up quickly by heat conduction. Finally it also melts and both are forming an sintering neck which is growing until both have flown together completely, which is the final stage. This cell model has been used to simulate the material behaviour in the powder bed during laser sintering, but it does not consider balling and clustering effects. These mechanisms have to be modelled additionally and might be subject of future investigations.startwettingcoalescencefinal stagelaser radiationr e y a lr e d w o pstandard material celld e r e t n i sr e y a lFig. 4:Model of the standard material cell4. FEM-SimulationBy use of the standard material cell, a simulation of the powder bed has been made by a FEM calculation for a laser pulse scanning. In fig.

4:Model of the standard material cell4. FEM-SimulationBy use of the standard material cell, a simulation of the powder bed has been made by a FEM calculation for a laser pulse scanning. In fig. 5a), the FEM model which represents the scanning of one powder layer on a sintered structure is illustrated. In the setup, the corner of a cube is radiated with a laser beam pulse of the power P for the exposure time texp.. The cube itself consists of a 50 µm thick powder layer on the surface and with a solid of the same material underneath. Originally, the laser beam has a round shape (diameter 0,5 mm) and gaussian energy distribution. The beam hits the cube with a quarter circle. In the calculations, the laser power P and exposure time texp. hav been varied. From the194exposure times, the referring scanning velocities v can roughly be calculated by the relation v = rB/texp.. In fig. 5b), the results for IN 718 are shown for three different parameter sets, e.g. P = 200 W/texp.

= 1300 °C – Tliq. = 1365 °CP = 170 W texp. = 10 ms (v = 50 mm/s) Tmax. = 1723 °C z (Tsol.) = 50 µm z (Tliq.) = 42 µm ∆zsol.-liq. = 8 µmP = 125 W texp. = 50 ms (v = 10 mm/s) Tmax. = 1591 °C z (Tsol.) = 50 µm z (Tliq.) = 36 µm ∆zsol.-liq. = 14 µmFig. 5:a) FEM model configuration and boundary conditions; b) process simulation for IN 718From the calculations in fig. 5b) it can be concluded, that scanning with low velocities and adapted low laser powers is more suited for laser sintering. With decreasing P and v, the maximum temperatures Tmax. become lower and the semisolid area ∆zsol.-liq. widens up. With the calculations the parameters of technical alloys can be derived. For IN 718, the parameter set P = 125 W/v = 10 mm/s is supposed to be the best from those in fig. 5. If the process should be speeded up by increasing the scanning velocity, a decrease of the layer thickness would be necessary.1955.

5. If the process should be speeded up by increasing the scanning velocity, a decrease of the layer thickness would be necessary.1955. SLS results of INCONEL 718On testing the nickel base alloy INCONEL 718 with SLS, sample cubes with the dimension 10x10x10 mm³ have been fabricated first. With these samples, the hardness and the metallographic structure have been analysed. Compared with conventional fabricated IN 718, a sensible decrease in hardness could be observed, see table 2. Until now, no final explanation has been found but it is assumed that the short process times of SLS do not allow mixed crystal formation, which causes a hardening effect [HAA 94]. The metallic structure of SLS is very fine and grain bounderies could hardly be observed, see fig. 6a) + b). The part has no visible porosity, but in some areas, failures, like pores, could be observed.

6a) + b). The part has no visible porosity, but in some areas, failures, like pores, could be observed. A hot isostatic pressing process was considered to eleminate these failures.IN 718 SLSIN 718 SLS + HIPIN 718 conventionalHardness (HV 0,2)Tensile strength (Rm)Fracture strain A5323 HV 0,2––268 HV 0,2750 N/mm²16,9 %538 HV 0,2 1430 N/mm² *) 21 % *)Table 2:Mechanical properties of differently fabricated IN 718 (*) values taken from [BEI 90])In the next step, tensile strength test bars have been produced by SLS, hiped for 4 hours at 1150 °C and at 1900 bar and were finally machined to net shape, see fig. 6c). The tensile test results are presented in table 2. The strength of the laser sintered structure of 750 N/mm² is quite low, compared with conventional IN 718. In the analysis of the test bars after the tensile tests, clusters of refractory oxides were found in the metallic structure, which itself is very ductile.

In the analysis of the test bars after the tensile tests, clusters of refractory oxides were found in the metallic structure, which itself is very ductile. It can also be seen that the deformation of the bars is caused by delamination. However, it is concluded, that oxydation during SLS and remaining, thin oxide skins on the SLS layers are the main reasons for the observed effects and has to be reduced intensely in future activities.50 µm50 µma) IN 718, after SLSb) conventional IN 718c) tensile strength test bars IN 718, SLS + HIP + machined to final shapeFig. 6:a) + b) micrographs of grinded and etched IN 718 structures; c) tensile strength test bars (length 65 mm, M10, diameter 6mm)1966. Summary and outlookOn a base test, the general coalescence behaviour of 2 metal grains was investigated. From the realised difference between semisolid and non-semisolid materials, it was concluded that process temperatures between Tsol. and Tliq.

From the realised difference between semisolid and non-semisolid materials, it was concluded that process temperatures between Tsol. and Tliq. within the actual powder layer are benefical to achieve satisfying process results. To permit temperature control, a process model for SLS was developed. It includes 4 submodels, e.g. the thermal, the metallurgical, chemical and sintering submodel. It can be used to predict or estimate process parameters, especially the combination of laser power, scanning velocity and layer thickness. The model has been applied for the nickel base alloy INCONELTM 718 by means of an FEM simulation and practical tests. It was possible to fabricate cubic sample parts for metallurgical analysis and bars for tensile strength testing. The mechanical strength of the material was not satisfying which is attributed to a high oxygen content in the material.Future work aims on machine and process modifications to reduce the oxygen content of laser sintered samples.

RECT SELECTIVE LASER SINTERING OF TOOL STEEL POWDERS TO HIGH DENSITY. PART A: EFFECTS OF LASER BEAM WIDTH AND SCAN STRATEGYC. Hauser, T.H.C. Childs, C.M. Taylor and M. Badrossamay School Of Mechanical Engineering, University of Leeds, UKS. Akhtar, C.S. Wright and M. Youseffi Engineering Materials Research Group, University of Bradford, Bradford, UKJ. Xie, P. Fox and W. O’Neill Department of Engineering, University of Liverpool, UKABSTRACTThis paper describes progress on the Direct Selective Laser Sintering of M2 and H13 tool steel powders, comparing this with previous and further observations on stainless steel powders. The distinguishing feature is the melting of single tracks and layers in deep powder beds. The paper focuses on changing characteristics of the melt pool (mass, volume, aspect ratio, stability) and laser-powder interactivity as the laser beam width, power and scan speed change.

The paper focuses on changing characteristics of the melt pool (mass, volume, aspect ratio, stability) and laser-powder interactivity as the laser beam width, power and scan speed change. It also compares the melt pool of neighbouring tracks during single layer construction. Simulations from a computer model to predict melt pool shape and dimension show reasonable agreement with experimental results at low scan speeds (0.5mm/s). But unexpected increases in melt depth above 1.0mm/s have been observed, suggesting higher values and more variability in laser absorptivity than expected, even approaching 1.0 for the CO2 laser radiation used in this work.INTRODUCTIONSelective melting of a metal alloy powder bed, aimed at direct production of one-off components and series tooling, is currently at the forefront of Selective Laser Sintering (SLS) research activities.

Recent advances have used specially blended alloys [1] and commercially available powders, including stainless steels [2-4] and tools steels [5,6]. However, reported part densities produced from these powders varies considerably, from just above the tap density of the virgin powder [4,6] to densities exceeding 99% of theoretical density [2,3]. Research in [4] has highlighted that understanding the behavioural differences and dimensional variations of the melt pool and rastering melt front as laser powers and scan speeds change allows for process refinement leading to improvements in the quality of melted layers. This paper extends [4] by also investigating the effects of laser beam width and powder material and how this information might be used to improve the quality of multiple layer components.

This paper extends [4] by also investigating the effects of laser beam width and powder material and how this information might be used to improve the quality of multiple layer components. The microstructure and density of tracks, layers and multiple layers produced in this work are examined in Part B [7].Running in parallel with the experimental activity is the continual development of a thermally-based finite element model which simulates the SLS process, being the subject of previous papers [8-10]. In this work the model has been used to generate plots of melt track cross-sections which are compared, geometrically, with similar images obtained from the experimental results.

In this work the model has been used to generate plots of melt track cross-sections which are compared, geometrically, with similar images obtained from the experimental results. It is this comparison between simulation and experiment which leads to conclusions of highly variable absorptivity in the powder beds.644THE FINITE ELEMENT MODEL AND ITS MATERIAL ASSUMPTIONSThe finite element thermal, powder melting and densification model has been described before, both in its 2-D [8,9] and 3-D [10] versions. It has been well-validated experimentally and quantitatively for the SLS of polymers [8,9].

It has been well-validated experimentally and quantitatively for the SLS of polymers [8,9]. For metals there are more severe issues of selecting correct material data than there are for polymers, when it comes to comparing simulations with experimental data, because of the larger temperature range to which the powder is exposed during processing, the larger difference between the powder bed and solid metal thermal conductivity and, particularly with the use of CO2 laser radiation, the lower and possibly more variable absorption of the radiation into the bed. In this paper handbook sources have been used for a metal solid's variation of thermal conductivity K and specific heat C with temperature, over the range 20(cid:113)C to 800(cid:113)C. A linear variation has been fitted and assumed to apply up to the melting temperature of the metal.

A linear variation has been fitted and assumed to apply up to the melting temperature of the metal. With T as absolute temperatureK = Ko + KvT; C = Co + CvTLatent heats of melting, L, have been obtained by a rule of mixtures from the latent heats of the alloys' constituent elements. The conductivity Kp and absorptivity (cid:68) of powder beds at room temperature have been calculated from experiments in which the surface of a bed was exposed to incident non-scanning radiation for a short period of time and the subsequent sub-surface temperature variation in the bed measured by thermocouples. The variation with depth of time for maximum temperature to occur was used to calculate the bed's thermal diffusivity - and Kp obtained from knowledge of the bed's density and material specific heat, and (cid:68) was obtained from the variation of maximum temperature rise with depth, both in the same theoretical way as described and used in [11].

Kp is assumed to vary with temperature in the same way as K.Estimates of track size, based on (cid:68) measured as just described, in many conditions grossly underestimate experimental observations, as will be seen later. Measured track mass per unit length, mL, has been used to obtain an alternative estimate of absorptivity, (cid:68)*. Equation 2 as an equality is an expression for the minimum value of (cid:68)*, based on a heat balance. (cid:68)*(P/U) is the absorbed energy per unit track length, with P the laser power and U the scan speed while mL[Cav.(Tm - To) + L] is the energy to melt unit track length, where Cav. is the average specific heat from ambient bed temperature To to the metal's melting temperature Tm.(cid:68)(cid:116)TCUPm (cid:62) (cid:64) (cid:11) (cid:12)(cid:62)(cid:11)Lav.M(cid:16)T o(cid:12)(cid:14)(cid:64)LEXPERIMENTATIONMaterials.

Three types of gas atomised powder have been used in this investigation: M2 (1C-4.15Cr-6.4W-5Mo-2V-bal.Fe) and H13 (0.4C-5Cr-1.3Mo-1.05V-bal.Fe) tool steels and 314S HC stainless steel (0.4C-20Ni-25Cr-1Si-bal.Fe). All powders were obtained from Osprey Metals Ltd, UK and were supplied with a particle size distribution of -150 + 75(cid:80)m. No heat treatments, additives or fluxes or powder pre-heating was used in this work. A particle size distribution of -75 + 38(cid:80)m was also investigated for examination in Part B [7] of this paper.645(1)(2)Equipment. SLS was performed in a research machine (Figure 1). A 250W continuous wave CO2 laser beam with a spot diameter adjustable between 0.55mm and 1.1mm at the focal length was scanned by galvanometer controlled mirrors on to the powder surface inside a controlled environment process chamber. A window in the chamber roof allowed the beam to enter. In the sealed chamber there were two build areas.

A window in the chamber roof allowed the beam to enter. In the sealed chamber there were two build areas. For multiple layer experiments the build area (70mm in diameter) was at the base of the chamber and on top of the piston assembly which could be moved up and down to control layer depth (accurate to 0.001mm). For single track and layer experiments the build tray (120mm x 150mm x 5mm deep) was used to maximise productivity. A hopper, driven by a stepper motor, deposited and spread fresh powder layers in one operation.Figure 1. Controlled environment process chamber.Melting Experiments. All experiments were carried out in an argon environment (obtained from 99.9% purity bottled argon). The chamber was evacuated to a rough vacuum followed by a one minute purge back to atmospheric pressure. This procedure was repeated twice before being followed by a 15 minute pre-sinter purge.

The chamber was evacuated to a rough vacuum followed by a one minute purge back to atmospheric pressure. This procedure was repeated twice before being followed by a 15 minute pre-sinter purge. During processing the flow rate of argon through the chamber was 3 litres/min and the net pressure was kept at 30mbar above current atmospheric pressure to ensure no air leakage into the chamber. Melting of all powder was carried out in a deep powder layer with a minimum first layer depth of 5.0mm. All scanning conditions for single tracks, layers and multiple layers are reported in Table 1. The experimental conditions were the same for all three powders and two beam widths of 0.55 and 1.1mm were used. The length of the single tracks was nominally 30mm, the size of the layers was nominally 15mm x 12mm and the multiple layers consisted of 12 and 25 layers.

The length of the single tracks was nominally 30mm, the size of the layers was nominally 15mm x 12mm and the multiple layers consisted of 12 and 25 layers. The multi-layer results will be reported elsewhere, some in Part B [7].Table 1: Experimental Conditions.ExperimentSingle tracksSingle Layer Multiple layersLaser Power (W)32, 58, 77, 110, 143, 170 58, 77, 110, 143 143Scan Speed (mm/s)0.5, 1 - 120.5, 1, 3, 5, 8, 10 5, 8, 10Scan Spacing (% of beam width)n/a25, 50, 100 25, 50Layer Thickness (mm) n/an/a 0.4646Simulations. Simulations of the single track experiments were carried out for all three powder materials. Results for a power of 77W, a beam diameter of 1.1 mm and for scan speeds (mm/s) of 0.5, 1, 2, 4 and 8 are presented here. Material properties used are given in the results section. GSharp V3.1 (AVS Inc) was used to produce 2D sectional views of modelled tracks, with a superimposed grid for dimensional analysis.

Material properties used are given in the results section. GSharp V3.1 (AVS Inc) was used to produce 2D sectional views of modelled tracks, with a superimposed grid for dimensional analysis. Image-Pro Plus V4 (Media Cybernetics) was used to evaluate the cross sectional solid area and take measurements from the images.Further Experiments. Measurements of powder bed diffusivity and absorptivity in the manner of [11] were carried out in air, with a CO2 laser beam of power up to 10W, defocused to a diameter of 4 mm at the powder bed surface. For these experiments, powder to a depth of 30 mm was placed in a cylindrical container of diameter 60 mm, with a thermocouple placed 10 mm beneath the surface coaxially with the laser beam.Video and series of still photographs were taken of single tracks as they formed, to gain insight into powder movement during processing.

These were taken through the viewing window (Figure 1) using a hand-held digital camera behind a MIG welding glass viewing filter.RESULTSProcess Maps. Single track process maps that detail heating and melting behaviour of the metal powders as scan speed and laser power change were produced first to highlight areas where the melt pool remained continuous i.e. did not fragment or ball. A fragmented melt pool has been shown in [4] to reduce the surface quality and density of layers. Process maps for M2 and H13 are given in Part B [7]. Figure 2 is for the 314S powder processed using a spot size of 1.1mm. The map is a refinement of the same process map originally presented in [12]. It shows several different qualitative regimes, common to all maps including those from different laser beam diameters. These are regions of continuous flattened tracks, continuous rounded tracks, melted and balled tracks, not fully melted and unmelted regions.

These are regions of continuous flattened tracks, continuous rounded tracks, melted and balled tracks, not fully melted and unmelted regions. Continuous tracks were removed from the bed and sectioned. Observations from all three powders, from P = 50 to 150 W and scan speeds up to 10 mm/s are considered in more detail next.Figure 2. 314S single track process map produced using a laser spot size of 1.1mm.647Track cross-section observations. Figure 3 shows track cross-sections for all three powders at powers of 77, 110 and 143 W. The boundary between the flattened and rounded tracks is added as the solid line. It is noticeable that track cross-section area increases with increasing scan speed as that boundary is crossed. The rounded tracks all have a much larger area than expected from the level of power absorption deduced from the embedded thermocouple tests to measure that (see later). Some qualitative observations may be relevant to this.

Some qualitative observations may be relevant to this. In the rounded track regime (the whole of the region enclosed by the dashed line in Figure 2), the tracks sink well into the powder bed. Powder to either side of a track collapses into it, leaving a trench surrounding the track. At low scan speeds (~1-2mm/s - Figure 3) track profiles are quite rounded, with only a few groups of satellite particles visible around the lower fringe of the solidified track. However, at higher speeds there are large clusters of sintered bonded particles at the lower surface, giving rise to 'bell' shaped track cross sections. Perhaps these differences arise from the time available for metal in the melt to flow - longer at low than higher scan speeds.Too fragile to handleFigure 3. Track cross-sections for M2, H13 and 314S powders, beam diameter 1.1 mm.648Track masses per unit length.

Track cross-sections for M2, H13 and 314S powders, beam diameter 1.1 mm.648Track masses per unit length. Figure 3 clearly shows increase of track cross-section with increase of scan speed at the boundary between flattened and rounded tracks. But the exact areas vary from place to place along a track. To obtain an average measure of size, tracks were removed from their powder beds and weighed. Figure 4 plots mass per unit length against speed for all three powders, for P = 77W. The two regimes, low speed (LS) and high speed (HS), are clear.Figure 4. Mass per unit length dependence on scan speed, P = 77W, beam diameter 1.1 mm.Figure 5 presents an alternative view, plotting mass per unit length against energy density (P/U), for two of the powders (as examples), but for all powers and speeds. In each case, data fall between upper and lower limits. All data for U > 4 mm/s fall on the upper limit, labelled HS. The lower limit labelled LS corresponds to U = 0.5 mm/s.

In each case, data fall between upper and lower limits. All data for U > 4 mm/s fall on the upper limit, labelled HS. The lower limit labelled LS corresponds to U = 0.5 mm/s. The HS and LS slopes for all materials, and experiments with both laser beam widths, are collected in Table 2, in units of milligram/J. According to equation 2, a constant slope mL/(P/U) corresponds to a constant 'effective' absorptivity, (cid:68)*, obtained from the equality of equation 2. Table 3 gives approximate values of (cid:68)* derived from Table 2 on the basis that Cav. = 700 J/kg and L = 280 kJ/kg - but see the next section for a more detailed consideration of absorptivity. The HS limits vary with beam diameter.H13H13Figure 5: Mass per unit length of 314S and H13 single tracks plotted against energy density for a 1.1mm beam width for 32 < P(W) < 170 and 0.5 < U(mm/s) < 12.649Table 2: Boundary gradients calculated from mass/length vs.

For this purpose, track area was defined as that within the simulated density boundary contour of 7,500 kg/m3. Table 5 summarises results for P = 77 W in the form of mL/(P/U), for comparison with the observations of Figure 5 / Table 2. The italic and bold entries are data obtained respectively by use of (cid:68)1 and (cid:68)2 in the simulations, to be compared with the LS and HS boundary values in Table 2. The simulations follow the same qualitative characteristic as the experiments. However, in the LS range, the data (with the exception of that for M2) range from 0.06 to 0.08, compared to 0.07 to 0.13 in Table 2: for H13, the simulation values are 6% higher than experiments; for 314S they are 36% lower. In the HS range, the data are approximately constant for each material, in line with experiment: comparing Tables 5 and 2, the simulation data are 63 (cid:114) 5 % of the experimental ones, for all three materials.Table 5.

mL,simulated / (P/U) (mg/J) - P = 77W, beam diameter = 1.1 mm.MaterialU (mm/s)M2 H13 314S0.5 1.0 2.0 4.0 8.0 0.15 0.47 0.51 0.48 0.48 0.07 0.08 0.48 0.46 0.41 0.06 0.07 0.08 0.28 0.27Alternatively, simulated results may be compared to the equivalent point experimental data in Figure 4. Table 6 records the ratio of the simulated to experimental values of mass per unit length. The italicised and bold scripts are obtained with (cid:68)1 and (cid:68)2, as in Table 5.Table 6. (mL,simulated / mL,experiment), P = 77W, beam diameter = 1.1 mm. MaterialU (mm/s)M2 H13 314S0.5 1.0 2.0 4.0 8.0 1.54 0.83 0.89 0.58 0.57 1.03 0.71 1.03 0.68 0.66 0.66 0.53 0.37 0.64 0.89The simulations and experiments differ from each other at both the LS and HS extremes, but in different ways. In the LS conditions, some simulations under- and others over-estimate the experimental observations, although on average they underestimate, and the more so the higher the scan speed.

In the LS conditions, some simulations under- and others over-estimate the experimental observations, although on average they underestimate, and the more so the higher the scan speed. It is probable that the main issue here is uncertainty in what are correct material data to enter into the simulation, particularly the value of (cid:68)1. In HS conditions, however, the simulations consistently predict track mass per unit length of around 65% of the observed values. This is despite, for the tool steels, absorptivity of 1.0 being assumed.There are two possible explanation. The tracks may contain a substantial amount (about one third of their mass) of unmelted powder, e.g. attached round their perimeters. Evidence to support this possibility comes from two directions: a) the simulated track masses reported here on the basis of the track boundary being the 7,500 kg/m3 boundary contour may be increased by c.

35% by expanding that boundary to include everywhere where any densification at all occurs; b) the experimental observations (Figure 3) frequently show irregular 'skirts' at the track bases (e.g. P = 77W, U = 2 to 10 mm/s for M2). Figure 7 is an enlarged view of the 4 mm/s track section. It shows the skirt occupying c. 30% of the track area. This is typical.651Or the model's assumption of heat transfer normal to the surface only by conduction is wrong. Associated with movements in the bed leading to cylindrical tracks, there may be powder motions that lead to more rapid movement of the melt front into the bed than by conduction. Powder movements are considered more next. Whether an absorption of 1.0 is realistic for CO2 laser radiation incident on to a steel powder bed is also addressed.Figure 7. M2 track cross-section enlarged from Figure 3, P = 77W, U = 4 mm/s.Track photography.

M2 track cross-section enlarged from Figure 3, P = 77W, U = 4 mm/s.Track photography. Viewing rounded track formation by eye, through the viewing window (Figure 1), revealed periodic melting of powder ahead of the main track. This has been recorded by video-filming through a MIG welding glass filter (actual tracks were bright orange, almost white, directly under the laser beam; without a filter clear images could not be obtained). Figure 8 is an example from filming a M2 track (P = 100 W, U = 1 mm/s).Figure 8. Video stills from track filming for M2: P = 110W, U = 1 mm/s, beam diameter 1.1 mm.652The first frame, at 0.01s, has marked on it the position of the axis of the incident laser beam, judged by where the melted track was brightest. 0.05s later, melting (marked 'secondary melting') is seen ahead of the track. This grows over the next 0.3s and finally flips (too fast to catch on film) into contact with the main track. The process starts again.

This grows over the next 0.3s and finally flips (too fast to catch on film) into contact with the main track. The process starts again. In summary, tracks do not grow in length steadily but as a series of steps. In this example, the periodic time is about 0.4s. This behaviour is seen qualitatively with all rounded tracks up to a scan speed of 4 mm/s, though filming has been only at 1 mm/s. As speed increases it becomes more difficult to resolve what is happening. A possible explanation of what is happening is illustrated by the sketch in the bottom right corner of Figure 8. Because the melt track is below the powder bed surface and has a steeply sloping nose, it is possible for the incident laser beam to be partly reflected from it on to the end wall of the powder trench in which the track sits. This double absorption opportunity will certainly lead to an increased (cid:68), but whether up to 1.0 is not clear.Single layer formation.

This double absorption opportunity will certainly lead to an increased (cid:68), but whether up to 1.0 is not clear.Single layer formation. Single layers have been created, based on understanding from the single track experiments. Two examples are discussed here (Figure 9), of M2 with P = 110W and U from 0.5 to 8 mm/s, at various scan spacings; and of H13 at P = 110W and U = 4 mm/s. First,Figure 9. Single layers, with scan spacing as a % of beam diameter; (a) 50%, (b) 100%, (c) 25%.the surface morphology of layers produced using M2 powder changed from flat to undulating and globular at the (P, U) condition where single tracks changed from flattened to rounded (Figure 3). Improvements in layer quality occurred with a scan spacing of 25% at higher speeds (8.0mm/s), but at lower speeds (1.0-2.0mm/s) the first two/three scan tracks within the layer were often globular and irregular.

At a scan spacing of 100%, layers produced at 8mm/s had an undulating surface morphology and the fusion bond between tracks was usually weak, causing areas of porosity. If a scan spacing, as a percentage of single track width is considered instead of as a % of beam diameter, globular layers occur at values lower that ~30% (as in the case for Figure 9a; 1-4mm/s and b, 2mm/s) and undulating and low porosity layers occur at values greater653than ~60%. Layers with improved surface quality shown in Figure 9a, 8mm/s and b, 4mm/s have a scan spacing relative to the melt track width of 40% and 36% respectively. The only exception to this rule occurred in layers produced from flattened tracks (0.5mm/s) where surface quality had less dependence on scan spacing.Figure 9c makes a different point. Its general condition (P = 110W, U = 4 mm/s) is in the rounded single track regime (Figure 3). The left panel shows the first track indeed to be rounded, but subsequent ones are not.

Its general condition (P = 110W, U = 4 mm/s) is in the rounded single track regime (Figure 3). The left panel shows the first track indeed to be rounded, but subsequent ones are not. Rounding of tracks is suppressed in layer formation with H13 powder more so than with M2 powder. How pre-existing tracks modify the absorptivity and heat capacity / conductivity environment of a currently melted track is currently under investigation. As a practical point, modifying the first scan power and speed (right panel of Figure 9c) to produce a flattened track at the start of each layer can improve the overall layer morphology.SUMMARY AND CONCLUSIONObservations of the formation mainly of single tracks, but also layers, of three different steels (M2, H13, 314S) in deep powder beds are reported. In the absence of a constraining solid substrate, tracks may form with a flattened section, a rounded section that sinks into the bed or as broken or balled segments.

In the absence of a constraining solid substrate, tracks may form with a flattened section, a rounded section that sinks into the bed or as broken or balled segments. The different laser power / scan speed conditions in which these occur are mapped and are similar for all three steels. A common observation is that there exists a scan speed range where melted mass increases with increased scan speed. This is not expected from constant power heating models. The speed range is the same as that in which single tracks change from a flattened to a rounded section with increasing speed. Track masses per unit length predicted by a thermal finite element model are in good agreement with experiment when flattened tracks are formed, using thermal absorption coefficients obtained independently, from measurements of temperature / time variations beneath a powder bed surface. To gain agreement when rounded tracks are formed, absorption coefficients close to 1.0 are required.

To gain agreement when rounded tracks are formed, absorption coefficients close to 1.0 are required. This has led to the video-filming of track formation that has revealed that rounded tracks do not form in a steady manner, but in a way that can enhance laser power absorption, whether up to 100% is another question. Another factor that leads to more track mass than is calculated - from the assumption that all of a track is melted matter - is the possible inclusion of unmelted or partially melted material round the bottom of a track.A remaining question to be answered is what causes the change from a flattened to a rounded track cross-section. Even without answering it, observations of single track quality can be used to plan the processing conditions for single layers.

Even without answering it, observations of single track quality can be used to plan the processing conditions for single layers. Here differences of detail occur between one tool steel and another (M2 and H13) with respect to the influence of laser power and scan speed on whether flattened or rounded multi-tracks are formed.ACKNOWLEDGEMENTThis work was carried out at the University of Leeds as part of a UK EPSRC funded project, ref. no. GR/R32239, linked to two other projects at the Universities of Bradford and Liverpool.654REFERENCES[1] M. Wohlert, S. Das, J. J. Beaman, D. L. Bourell (1999). Direct laser fabrication of high performance metal components via SLS/HIP. Proc.10th Solid Freeform Fabrication Symposium, University of Texas, Austin, pp. 281-288.[2] W. Meiners, C. Over, K. Wissenbach, R. Poprawe (1999). Direct generation of metal parts and tools by Selective Laser Powder Remelting (SLPR).

RECT SELECTIVE LASER SINTERING OF TOOL STEEL POWDERS TO HIGH DENSITY: PART B – THE EFFECT ON MICROSTRUCTURAL EVOLUTIONS Akhtar*, CS Wright* and M Youseffi* *Engineering Materials Research Group, University of Bradford, Bradford, UKC Hauser#, THC Childs#, CM Taylor# and M Baddrossamay# #School Of Mechanical Engineering, University of Leeds, UKJ Xie+ and P Fox+ and W O’Neill+ +Department of Engineering, University of Liverpool, UKAbstractThis paper describes recent progress on the Direct Selective Laser Sintering of M2 [Fe-6W-5Mo-4Cr-2V-0.8C] high speed steel (HSS) and H13 [Fe-5Cr-1V-1Si-1.5Mo-0.4C] tool steel powders. Part B will focus on the microstructural evolution of laser scanned powder beds. It has been found that H13 powders are more amenable to Direct Selective Laser Sintering than M2 powders. Densities up to 90% are possible with H13 powder compared with 70% for M2.

It has been found that H13 powders are more amenable to Direct Selective Laser Sintering than M2 powders. Densities up to 90% are possible with H13 powder compared with 70% for M2. The relationship between alloy composition, microstructure, post-scanned density and scan conditions will be discussed for single track, single layer and multi-layer constructions.IntroductionSelective laser sintering (SLS) is one of the more important solid freeform manufacturingprocesses developed in the last 10-15 years [1]. One of the main applications of SLS is for the manufacture of prototype or low volume production tooling using steel powders. Commercial systems are available which generate parts from either polymer coated steel powders [DTM RapidTool 2.0TM] or by processing a mixture of metal powders that contains a low melting point component [EOS DirectToolTM]. The need to post process, i.e.

The need to post process, i.e. infiltrate with bronze, or the reliance on a low melting point liquid, leads to components with poor mechanical and tribolological properties. This has led to a number of studies aimed to investigating Direct Selective Laser Sintering (DSLS) pre-alloyed metal powders with the aim of eliminating the above disadvantages. These studies generally made use of commercially available powder stainless steel [2, 3], and tool steels [4, 5]. Recently Dewidar [6] has reported studies on an experimental Mo-based High Speed Steel, a composition developed by Wright et .al.[7], specifically as a “process-friendly” grade for processing by conventional die pressing and vacuum sintering. Whilst it is possible to process Stainless Steel powders to densities of +99% [3], less success has so far been achieved with tool steels.

Whilst it is possible to process Stainless Steel powders to densities of +99% [3], less success has so far been achieved with tool steels. This paper presents results of investigations into the melting / densification behaviour of tool steel powders during DSLS which has the aim of identifying the factors which control the ability to fabricate fully dense, multi-layer components.656Experimental TechniquesThree types of powder were used in this investigation. Annealed water atomised M2(WA-M2), gas atomised M2 (GA-M2) and gas atomised H13 (GA-H13). The water atomised M2 was sourced from Powdrex, UK and the gas-atomised powders were obtained from Osprey Metals Ltd, UK.

The water atomised M2 was sourced from Powdrex, UK and the gas-atomised powders were obtained from Osprey Metals Ltd, UK. The various size fractions used, along with Apparent Density and flowability (Hall Flowmeter) data are given in Table 1.Powder TypeWA-M2GA-M2GA-H13Size Fraction ((cid:80)m) <150/+75 <75/+38 <38 150/+75 <75/+38 <38< <150/+75 <22Apparent Density (gcm-3) 2.14 2.18 2.60 3.97 ND 4.31 4.17 3.41Flowability (s/50g) 55.4 59.8 No flow 23.7 ND No flow 20.2 no flowTable 1. Powder Properties of materials investigatedSelective laser sintering was performed on specially constructed machines at theUniversities of Leeds and Liverpool. The Leeds machine includes a 250W continuous wave (CW) CO2 laser. Preliminary studies were conducted at a Beam size of 1.1mm diameter. This was subsequently reduced to 0.55mm diameter.

The Leeds machine includes a 250W continuous wave (CW) CO2 laser. Preliminary studies were conducted at a Beam size of 1.1mm diameter. This was subsequently reduced to 0.55mm diameter. Galvanometer controlled mirrors direct the laser beam within a 70.0mm diameter build area which is housed in a 300 cm3 (L = 460mm, H = 260mm, and D = 250mm) chamber capable of sustaining a variety of atmosphere conditions including an absolute pressure of 10mbar. Details of this machine are given in Reference 2. In this study, the atmosphere inside the chamber was high purity argon (bottled argon at 99.9% purity). The chamber was evacuated to 30 mbar followed by a 15 min pre-sinter purge with argon at atmospheric pressure. During sintering the gas pressure inside the chamber was 30mbar above atmospheric. The Liverpool system, described by Morgan, et. al. [3], consists of a Rofin Sinar 90W, flash lamp pumped Q-Swiched Nd : YAG laser.

The Liverpool system, described by Morgan, et. al. [3], consists of a Rofin Sinar 90W, flash lamp pumped Q-Swiched Nd : YAG laser. An analogue galvanometer-scanning head attains beam position over an area of 80 x 80 mm2. Pulse repletion rates are in the range 0 – 60kHz and minimum beam diameter is 80µm. The shroud gas was nitrogenThe two systems employed different approaches to placing layers of powder. In the Leeds machine, the initial powder layer comprised loose powder to a depth of 5 mm contained in a mild steel tray. This layer was levelled with a blade to ensure a flat powder surface. Subsequent layers were deposited using a hopper system. The layer depth was 0.4 mm. In the Liverpool machine mild steel sheet was used as a substrate. In this system, the powder delivery system, which employs a counter rotating roller system [3], was optimised for an even 100(cid:80)m coating of powder to be layered for every build layer.

In this system, the powder delivery system, which employs a counter rotating roller system [3], was optimised for an even 100(cid:80)m coating of powder to be layered for every build layer. Whilst all size fractions studied could be processed on the Leeds machine, the Liverpool machine was restricted to processing just the <38(cid:80)m and <22(cid:80)m fractions.657After sintering samples were prepared for metallographic examination by mounting andgrinding on SiC paper to 1200 grit, then polished on cloths impregnated with 6 and 1(cid:80)m diamond. Polishing was completed using 0.05(cid:80)m (cid:74)-alumina. Samples were studied in the etched and un-etched conditions. M2 samples were etched in 5% Nital whilst 10% Nital was used for the H13 specimens.

Samples were studied in the etched and un-etched conditions. M2 samples were etched in 5% Nital whilst 10% Nital was used for the H13 specimens. Samples were examined using both optical and scanning electron microscopy techniques.Isopleths were calculated for the M2 and H13 systems using ThermoCalcTM[www.thermocalc.se] software and the SGTE SSOL database in order aid with both the selection of post-processing conditions and interpretation of microstructures, Figure 1.(a)(b)Figure 1. Calculated Isopleths for (a) Fe-6W-5Mo-4Cr-2V-C, and (b) Fe-5Cr-1V-1Si-1.5Mo-C systemsResults and DiscussionM2 PowdersTypical SEM micrographs of the “as-received” powders are given in Figure 2.100µm50µm100µm(a)(b)(c)Figure 2. SEM micrographs of M2 powders. (a) WA-M2 <150µm, (b) GA-M2 <38µm,(c) GA-M2 <150/+75µm size fractions.The water-atomised powder was composed of highly irregular particles and, as such, istypical of this method of powder production.

As expected the gas-atomised powders were composed of spherical particles. However, whilst the >38(cid:80)m GA-M2 size fraction consisted of658spherical particles with smooth surfaces [Figure 2(b)] the particle surfaces of the coarser fractions were much rougher due to the presence of many smaller satellites formed as a result of collisions during atomisation, Figure 2(c).20µm50µm(a)(b)Figure 3 Microstructures of M2 powders. (a) Gas atomised <150/+75. Note the finecarbide network. (b) Water atomised <38(cid:80)m fraction. Note fine discrete carbides.At the outset of this work two strategies were envisaged for the manufacture of highdensity, multi-layer components;(a) Partial melting of successive powder layers forming sufficient liquid to bond both powders and layers together, followed by a post processing sintering step. High Speed steels are one of the few alloys systems that are commercially sintered to full density (i.e.

High Speed steels are one of the few alloys systems that are commercially sintered to full density (i.e. +99% relative density) via a supersolidus liquid phase sintering mechanism [7]. The aim of the post processing step would be to vacuum sinter the laser sintered performs at a temperature of ~1260oC, the optimum sintering temperature for conventional die-pressed M2 powders [10], initiate controlled re-melting, inducing densification through particle rearrangement and solution – re-precipitation mechanisms [7]. The challenge was to produce intact laser sintered performs which, in order to avoid distortion during post- processing sintering, contained uniform distributions of porosity(b) Complete re-melting of powder layers.The heating and melting behaviour of each batch of M2 powder was systematicallystudied using the CO2 laser SLS machine at Leeds in order to identify appropriate processing conditions to achieve the two processing strategies.

Single tracks were produced for a wide range of Laser power and scanning conditions. In common with previous investigations [2, 6] the results for M2 are presented as process maps, Figure 4.659(a)(b)Figure 4. CW CO2 laser process maps for gas-atomised M2 <150/+75µm fraction (a)1.1mm and (b) 0.55mm spot size.A feature of the process maps was that the size of the regimes and the positions of theboundaries defining each regime were comparable, irrespective of whether the powder was gas or water atomised and size fraction. Reducing spot size had a only small effect on increasing the size of the continuous melting regime, e.g. Figures 4(a) and 4(b).

Figures 4(a) and 4(b). Attempts were made to fabricate single and multi-layer blocks (12 x 12 x 12 mm3) using conditions located within the partial melting and continuous melting regimes using the CO2 laser SLS machine, Table 2.Size Fraction (µm)<38water atomised <38gas atomised <75/+38water atomised <150/+75gas atomised <150/+75gas atomised <75/+38gas atomised <75/+38water atomisedBeam Overlap (%) 25 25 25 25 25 25,75 25,50,75Power (W)60 50,100 50,100 50, 60 130, 135 130 130Scan Speed (mms-1) 60, 100 70,130 70, 130 100,130 8 5,8 5,8,10Spot Size (mm)1.1 1.1 1.1 1.1 0.55 0.55 0.55CommentsSingle Layer Single Layer Single Layer Single Layer Multiple Layer Multiple Layer Multiple LayerTable 2.Conditions used to process M2 powders using CW CO2 LaserIrrespective of powder type or particle size fraction, attempts to produce multi-layer blocks were unsuccessful using conditions located within the partial melting regime were unsuccessful.

The sub 38µm water and gas atomised size fractions were difficult to spread, layers tended to form660furrows. However the major issue with processing within the partial melting regime was that at the conditions studied, insufficient melting took place to bond the next layer on to the previous layer, making it impossible to build intact blocks. Thus it was not possible to pursue fabrication of multi-layer components for M2 via strategy (a). Multi-layer blocks made from 25 individual layers were fabricated using conditions within the continuous melting regime. Unfortunately, the resultant blocks were very porous.Using the Liverpool SLS system, it was possible to scan at up to 500 mms-1 and operate either in pulsed or continuous modes. Initial experiments were performed using water atomised M2 at scan speeds of 10 –500mms-1, laser current 16-20A, Spot size = 0.1 mm, Beam overlap of 25-80%, pulse repetition frequency of 0kHz. The layer thickness was 100µm.

The layer thickness was 100µm. Both single and 3 layer samples were produced. A feature of the scanned areas, though, was that they contained large amounts of un-melted powder. Removal of this by ultrasonic cleaning revealed that the beds were very porous, Figure 5(a). The porosity levels were unaffected by scanning conditions. This high level of porosity was attributed to the low apparent density of the water atomised powder arising from its highly irregular shape, Table 1, and the layer thickness. In previous studies with 316L stainless steel, operating in pulsing mode blocks with 100% relative density being obtained [4]. Consequently, additional experiments with M2 powders were carried out at layer thickness of 100 and 50µm, pulse repetition frequencies of 40 and 50kHz, scan speed = 500 mms-1, current = 20A, spot size = 100µm, beam overlap = 50%. Blocks 5 x 15 x 15mm3 were produced. Operating in pulse repetition mode resulted in an increase in density compared to continuous mode.

Blocks 5 x 15 x 15mm3 were produced. Operating in pulse repetition mode resulted in an increase in density compared to continuous mode. However densities were only 5.5 gcm-3, i.e. ~70% relative, and were independent of processing conditions, Figure 5(b).(a)(b)Figure 5. Micrographs of Nd : YAG laser scanned multi-layer M2 (a) <38µmWA-M2, 400 mms-1scan speed, 50% beam overlap, current 20A, spot size 0.1mm, (b) 500 mms-1scan speed, 50% beam overlap, current 20A, spot size 0.1mm, pulse rate 40kHz. (100µm layer thickness).Despite the higher apparent density of the <38µm gas atomised powder, Table 1, use of this powder had no beneficial effects on the density of laser scanned blocks, densities were similar to those obtained for <38µm water atomised M2 powder.661The microstructures of samples produced on the two systems were comparable and independent of starting powder.

Single layers produced using the CO2 laser operating within the partial melting regime contained regions of un-melted powder, Figure 6(a). The melted regions had a very fine cellular - dendritic structure (grain size <10µm) with fine eutectic networks at grain boundaries, Figure 6(b). Partially melted regions were also visible [Figure 6(c)].M2CFerriteM6C20µm10µm20µm(a)(b)(c)Figure 6 CW CO2 M2 single layers at 1.1mm spot size. (a) <38µm, 60W, 25% beam overlap, 100mms-1 Scan speed, (b) <38µm, 60W, beam overlap 25%, scan speed 70mms-1, (c) <38µm, 60W, beam overlap 25%, scan speed 100mms-1.The dark areas [Figure 6(c)] are believed to be ferrite, suggesting that these regions been reheated to within the L + (cid:534)+ (cid:302) or L +(cid:534) region of Figure 1(a). Two types of eutectic were observed. Given the small size of the eutectics it was not possible to identify them using X-ray microanalysis.

Two types of eutectic were observed. Given the small size of the eutectics it was not possible to identify them using X-ray microanalysis. However based on both grey levels in Back Scattered Electron imaging (Atomic Number contrast) and eutectic morphology, it is possible to identify the carbide components as M6C (Fe-W-Mo rich) and M2C (Mo-W-Fe –V rich) [8]. Irrespective of composition or cooling rate, M6C adopts a characteristic “herringbone” structure. The presence of M2C is not expected from Figure 1(a), instead V rich MC would be expected to form under equilibrium cooling. High cooling rates, such as those of laser processing techniques, favour M2C formation over MC [8]. M2C eutectic can adopt one of three morphologies, feathery, cellular or lamellar [8]. High cooling rate also favours the formation the cellular morphology observed in the SLS samples [8].

M2C eutectic can adopt one of three morphologies, feathery, cellular or lamellar [8]. High cooling rate also favours the formation the cellular morphology observed in the SLS samples [8]. Microstructures of single layer samples processed in the continuous melting regime were comparable, the only difference being a reduction in the amount of un-melted powder present.Examination of the multi –layer samples revealed that re-melting was taking place to adepth of 2-3 layers. For example in the blocks produced from 25 powder layers on the CO2 laser system only 13 layers were visible afterwards. In terms of eutectic type and morphology, presence of partially melted regions, the microstructures were identical to the single layer samples. However an important additional microstructural feature was the presence of coarse eutectic layers within the blocks, e.g.

However an important additional microstructural feature was the presence of coarse eutectic layers within the blocks, e.g. Figure 7(a), believed to be formed by epitaxial solidification at solid-liquid interfaces.66210µm50µm(a)(b)Figure 7 Micrographs of M2 CW CO2 laser multi-layer specimens (a) GA-M2<150µm<38µm, 60W, 25% beam overlap, 70mms-1 Scan speed, 0.55 mm spot size (b) Post processed sample – 1260oC for 60 min.Eutectic structures in High Speed Steels are considered to be detrimental to mechanical properties. Hoyle [9] demonstrated for conventional cast HSS that it was possible to refine the “as cast” eutectic structures by reheating to just above the solidus followed by slow cooling.

Hoyle [9] demonstrated for conventional cast HSS that it was possible to refine the “as cast” eutectic structures by reheating to just above the solidus followed by slow cooling. In a similar manner, the multi-layer blocks were subjected to a post processing step which involved heating in vacuum at 10oCmin-1 to 1260oC, the optimum Supersolidus Liquid Phase Sintering temperature for M2 powders [10], holding for 60min then furnace cooling at an average rate of ~2oCmin-1 to room temperature. Whilst this treatment did not produce any further densification, it resulted in the breakdown of the eutectic structures; networks being replaced by fine, discrete carbides with regions such has those of Figure 7(a) being replaced by coarser carbide clusters, Figure 7(b).

This is considered significant because even if M2 powders could be processed to ~100% density by DSLS, such regions would be potential points of weakness since it is known that in cast wrought and powder metallurgy processed HSS that carbide clusters are preferred sites for crack nucleation at low applied stresses under static and dynamic loading [11].H13 PowdersThe Gas atomised H13 powders were spherical with minimal satellites. The microstructure of “as-received” powders comprised cellular-dendritic grain structure with continuous grain boundary carbide films, Figure 8. The grain size decreased with particle size.10µm10µm(a)(b)Figure 8. Microstructure of gas atomised H13 <150/+75(cid:80)m size fractionFigure 9 shows the Process Map for <150/+75µm powder processed using a 0.55mm spot size on the CWCO2 system.663Figure 9.

Microstructure of gas atomised H13 <150/+75(cid:80)m size fractionFigure 9 shows the Process Map for <150/+75µm powder processed using a 0.55mm spot size on the CWCO2 system.663Figure 9. CW CO2 laser process maps for gas-atomised H13 <150/+75 µm fraction,0.55mm spot sizeComparing Figure 9 with Figure 4(b), it can be seen that the process map for H13 is comparable to those for M2 powders. Interestingly, processing H13 powders within the partial melting regime of Figure 9 was slightly more successful than with M2 in that it was possible to produce multi-layer blocks. The individual layers were ~1/3 thinner than for M2 processed in the same regime but these multi-layer blocks were very porous and fragile, Figure 10(a).Much more promising results were obtained with H13 compared to M2 in the Nd : YAG SLS.

Processing at condition of Pulse repetition frequency of 30kHz, scan speed 500mms-1, current 20A, spot size 100µm, beam overlap 50% give well-bonded, multi-layer blocks with a density of 90%, cf. Figure 10(b) and Figure 5(b).200µm(a) <150/+75µm size fraction(b) <22µm size fractionFigure 10. H13 multi-layer samples (a) CW C02 laser at 0.55mm Spot size size, 5mms-1 scan speed, 50% Scan spacing, 32W Power. (b) Nd : YAG laser at 500 mms- 1scan speed, 50% beam overlap, current 20A, spot size 0.1mm, pulse rate 30kHz. (100µm layer thickness)664Figure 11 shows typical microstructures produced for continuous melting using the Nd:YAG system and partial melting using the CO2 system.10µm30µm20µm(b) Figure 11. Micrographs of H13 powders scanned at same conditions as Figure 10.

Micrographs of H13 powders scanned at same conditions as Figure 10. (a) Nd : YAG laser [Pulsed Mode], (b) and (c) CW CO2 laser(a)(c)Processing using the Nd : YAG laser produced a fine, grained cellular – dendritic size with continuous grain boundary carbide films, Figure 11(a). The matrix was martensitic. Carbide depleted zones were observed [Figure 11(a)] and in some areas there was no bonding between layers. Due to the fineness of the microstructure, this carbide phase could not be identified using X-ray microanalysis. However, unlike M2, coarse carbide clusters were not observed at interfaces. Superficially the microstructures of the samples produced by processing in the partial melting regime of the CO2 laser were similar to those produced by Nd : YAG laser scanning. They appeared to consist of a cellular – dendritic grain structure, a martensitic matrix and continuous grain boundary carbide films, Figure 11(b) and (c).

They appeared to consist of a cellular – dendritic grain structure, a martensitic matrix and continuous grain boundary carbide films, Figure 11(b) and (c). Detailed observation using SEM revealed a more complex situation in that the network features were simply regions of differential contrast, they were crossed by austenite grain boundaries and by individual martensite laths, Figure 11(c). X-ray microanalysis indicated that these regions were slightly richer in Chromium than the grain interiors. This type of structure could arise due to reheating of a previous solidified layer during subsequent scanning of the top layer of powder. The carbides present in laser scanned H13 materials have yet to be positively identified but it can be inferred from Figure 1(b) that non-equilibrium solidification will result in the formation of Cr rich M7C3 carbide in preference to vanadium-rich MC predicted for equilibrium solidification.

M7C3 carbides readily dissolve in austenite at temperatures above ~950oC. Reheating for a short time, at a temperature below the solidus sufficient to dissolve the carbides, yet high enough to form austenite, followed by rapid cooling could, in principal, produce the microstructure seen in Figure 11(c). Carbon diffusion rates are much high in (cid:534)-Fe than Chromium diffusion rates and, if the time at temperature is insufficient for homogenisation, a coring effect will result, as indicated by X-ray microanalysis. Once the carbides have dissolved there will be no pinning effect and an austenite grain structure will form with grain boundaries cutting across the cored regions. The austenite will transform to martensite on cooling with individual laths passing through cored regions.

The austenite will transform to martensite on cooling with individual laths passing through cored regions. It is not yet understood why this structure was not observed in the Nd : YAG processed powders.The results presented here indicate, as with the previous studies of Nui and Chang [4, 5], that gas-atomised powders are more suited to the DSLS approach than water-atomised powders. Whether this is due to the higher oxygen content of water-atomised powders, as suggested by them or particle morphology have still to be established. For example, are the differences in665behaviour due to lower apparent density of water atomised producing less dense powder beds or is due to the bed powder flowability and how the powder interacts and flows into the moving melt pool. More interestingly, the current studies have shown that H13 powders are more amenable to DSLS than M2 powders.

More interestingly, the current studies have shown that H13 powders are more amenable to DSLS than M2 powders. Given that powder properties for the two compositions are comparable this would suggest that alloy composition has an important role to play, particularly since similar sized stainless steel powders can be processed to 100% density [3]. The reasons for this are the subject of ongoing investigations.Conclusions[1] M2 high speed steel powders are not amenable to processing to full density by a DSLS processing. The maximum density achieved in multi-layer samples, using pulsed Nd : YAG processing was only 70% relative.

The maximum density achieved in multi-layer samples, using pulsed Nd : YAG processing was only 70% relative. Much lower densities were obtained using CW CO2 and CW Nd : YAG processing operating in the same continuous melting regimes.[2] Gas-atomised H13 powders are amenable DSLS with densities of 90% relative being achieved using pulsed Nd : YAG processing operating in a regime of continuous melting[3] The reasons for the differences in processing behaviours of M2 and H13 powders of similar powders properties are, as yet, unclear.AcknowledgementsThe research reported in this paper is a joint project between the Universities of Bradford, Leeds and Liverpool, funded by the UK Engineering and Physical Sciences Research Council under Grant Number GR/R32222.References[1][2][3]KW Dalgarno and CS Wright; “Approaches to Processing Metal and Ceramics through the Laser Scanning of Powder Beds”, Society of Manufacturing Engineers, Technical Paper PE03-143, 2003.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDirect Selective Laser Synthesis of CuCrFeNiTiAl High Entropy Alloy from Elemental Powders through Selective Laser MeltingJoni Dhar1, Lazaro Lopez1, Shanshan Zhang1, Ben Xu2, Jasim Uddin1, Jianzhi Li11Manufacturing and Industrial Engineering, University of Texas Rio Grande Valley,Edinburg, TX 785392Mechanical Engineering, Mississippi State University, Mississippi State, MS, 39762AbstractThis study investigated the synthesis of CuCrFeNiTiAl high entropy alloy (HEA) from pure elements using selective laser melting (SLM). The objectives are to validate the feasibility of the HEA fabrication from elemental powder materials, and to examine the effect of various process conditions in SLM, such as laser power, point distance and laser exposure time, on the microstructures formed.

The as-built samples under high, medium and low energy densities were characterized by X-ray diffraction (XRD), and the microstructures were observed using scanning electron microscopy (SEM). The XRD results showed that five major crystal structure phases (hexagonal, monoclinic, orthorhombic, body-centered cubic and rhombohedral) were present in all samples. Fine-grained phases were noticed on the sample surface with non-uniform microstructural distribution. Such phases in high and low energy density samples were observed polygonal while round-shaped microstructures were observed for samples prepared under medium energy density conditions. Also, the grain size was proportional to energy levels of the fabrication process. Large size and clustered structures are prominent in samples produced under high energy density.Keywords: selective laser melting, high entropy alloy, microstructure, additive manufacturing1.

Large size and clustered structures are prominent in samples produced under high energy density.Keywords: selective laser melting, high entropy alloy, microstructure, additive manufacturing1. IntroductionHigh-entropy alloys (HEA) as a new family of alloys are composed of five or more principal elements in equimolar or near-equimolar concentrations [1, 2]. HEAs contain principal elements with concentration between 5 to 35 at %, and minor elements each below 5 at % [2]. By combining multiple elements in an alloy system, HEAs are showing a great possibility for designing novel materials with superior properties expected under extreme operating conditions (high temperature, high pressure, etc.) in many applications in the defense and energy industries such as nuclear reactors, missiles, jet engine, nozzle, turbine blade, compressor blade of aero engine, etc.The earliest attempts to design high entropy alloys are the research works of two research groups led by Prof. B.

B. Cantor and Prof. J.W. Yeh in 2004 [3, 4]. Both the groups investigated multicomponent alloy systems, which triggered the development of high entropy alloy. Yeh is the first who thoroughly described four core effects of high entropy alloys: high-entropy effect, lattice distortion, sluggish diffusion, and cocktail effect [5]. High-entropy effect tends to stabilize the HEA phase by minimizing the Gibbs free energy. Severe lattice distortion effect describes the displacement of lattice sites in HEA system from the perfect crystal lattice site. Sluggish diffusion958effect explains how the disorder of the multi-component HEA phase causes the local lattice potential energy to form deep traps that inhibit diffusion among the elements. Cocktail effect is the selection of the specific elements with specific properties that will contribute to the desired properties of the final high-entropy system.Following the above-mentioned pioneering works, researchers have studied many HEA systems.

CuCrFeCoNiAl [6-9] and its Cu free derivations [10, 11] are the most comprehensively studied alloy systems among all the HEA systems reported. One of the important derivatives is CuCrFeNiTiAl where Co is replaced by Ti. As titanium has the highest negative mixing enthalpy with rest of the elements, inclusion of titanium in HEA system causes different intermetallic phases to grow. These phases strengthen HEA with improved strength and hardness [12, 13]. Besides, the solid solution strengthening mechanism of titanium, due to its higher atomic radius, enhances strength of alloys greatly[13]. Therefore, CuCrFeNiTiAl HEA system have a high potential of creating HEAs with promising performance.With the growing interest in research on HEA, many HEA fabrication systems have been invented such as arc melting, plasma arc sintering, casting, etc. [14-18].

[14-18]. Unlike conventional manufacturing techniques, powder-bed additive manufacturing (AM), such as selective laser melting (SLM) process, has been investigated for HEA fabrication [19, 20]. Studies were reported in literature focusing on the effect of printing parameters (laser power, hatch distance, point distance, exposure time, laser focus diameter, scanning rotation angle, etc.) on the microstructure and properties of alloys made by SLM process [21-23]. However, the reported feedstock of these SLM fabricated HEAs are mostly based on pre-alloyed powders [24], which show disadvantages such as high cost, and lack of flexibility of material designs.The objective of this study is to demonstrate the feasibility of HEA fabrication from elemental powders via SLM process, and to examine the effect of different process parameters in SLM, including laser power, exposure time, and point distance, on the microstructures of the specimens.

A CuCrFeNiTiAl high entropy alloy system was synthesized directly from elemental powders under different energy densities. Then, the as-built specimens were characterized using X-ray diffraction (XRD) for crystal structures and observed under scanning electron microscopy (SEM).2. Experimental details2.1. MaterialsThe CuCrFeNiTiAl HEA powder was blended using 99.9 % pure (trace metal basis) Cu, Cr, Fe, Ni, Ti and Al elemental powders, supplied by Sigma-Aldrich Co. The nominal composition of the powder blend is given in Table 1. The elemental powders were measured and mixed in an oxygen-free sealed chamber filled with argon (Ar) atmosphere and then placed into a V-type blender rotating for 2 hours at 200 rpm.Table 1. Chemical composition of the powder blend.Element Weight, % Molar Ratio, %Al 2.50 5Cu 22.36 19Cr 18.20 19Fe 19.59 19Ni 20.61 19959Ti 16.72 192.2.

Chemical composition of the powder blend.Element Weight, % Molar Ratio, %Al 2.50 5Cu 22.36 19Cr 18.20 19Fe 19.59 19Ni 20.61 19959Ti 16.72 192.2. SLM FabricationA Renishaw AM 250 SLM system was used for fabricating the samples using the blended HEA powder. This system utilizes a high precision fiber laser with 1070 nm wavelength, 70 μm focal diameter and 200 W power. The laser scanned the powder surface during the SLM process in pulsed mode. The duration of heating by the laser at a single point is called exposure time and the distance between two consecutive points along a hatch line is called point distance. Hatch spacing is defined as the distance between two consecutive hatch lines. To save the powder materials and build time, a reduced build volume (RBV) system was developed, as shown in Figure 1. The build substrate is customized and machined using 316L stainless steel with a dimension of 50.8 mm × 50.8 mm.

The build substrate is customized and machined using 316L stainless steel with a dimension of 50.8 mm × 50.8 mm. To investigate the effect of different process conditions on the microstructure of the synthesized CuCrFeNiTiAl samples, a 5 × 5 test series was designed with different process parameter combinations. Each sample in the test series was designed as a thin square chip with the dimension of 6.4 mm × 6.4 mm × 0.84 mm. The size of substrate in the RBV system was considered during the selection of this dimension.Figure 1. Reduced build volume system used in this work.In this study, the process parameters were investigated, including laser power, exposure time, and point distance, as shown in Table 2. The hatch spacing and layer thickness were kept constant at 90 µm and 70 µm, respectively. A unidirectional meander build strategy was applied with an alternative rotation of 67° in the consecutive layers.

A unidirectional meander build strategy was applied with an alternative rotation of 67° in the consecutive layers. The prescribed parameters, such as laser beam size and laser focus offset were used the machine default setting. During the fabrication, the oxygen concentration inside the chamber was set 500 ppm (parts per million) and high purity argon gas with a line pressure of 1 bar was used as inert media. For facilitating proper melting of the first powder layer, the build substrate was preheated to 170 °C. When the process chamber was ready, the HEA powder blend was spread using the in-house developed RBV system.960Table 2.

When the process chamber was ready, the HEA powder blend was spread using the in-house developed RBV system.960Table 2. Experimental design of the 5 × 5 test series.Point Distance (µm) →60555045Power (W)↓ 100 125 150 175 200Exposure Time (µs)↓ 20 35 50 65 801A 2A 3A 4A 5A1B 2B 3B 4B 5B1C 2C 3C 4C 5C1D 2D 3D 4D 5DAccording to Gu and Shen [25], the energy density (ED) can be calculated in terms of laser power (P), scan speed (v), hatch spacing (h) and layer thickness (l). However, in the Renishaw AM250 system, the scan speed of the pulsed laser is determined by the ratio of point distance (d) and exposure time (t). Therefore, the energy density in this study is evaluated by Equation (1). In this study, three samples under high, medium and low energy densities were selected, marked as 1A, 3C and 5E in Figure 2. Per Equation (1), the energy densities for Samples 1A, 3C and 5E are 5.29 kJ/mm3, 23.81 kJ/mm3, and 63.49 kJ/mm3, respectively as shown in Table 3.

Per Equation (1), the energy densities for Samples 1A, 3C and 5E are 5.29 kJ/mm3, 23.81 kJ/mm3, and 63.49 kJ/mm3, respectively as shown in Table 3. ⋅ P t ⋅ ⋅ d h lED=Figure 2. (a) Detailed dimension of the test series, and (b) SLM fabricated samples.Table 3. Energy densities (kJ/mm3) of test series.Point Distance (µm) →60555045Power (W)↓ 100 125 150 175 200Exposure Time (µs)↓ 20 35 50 65 805.29 11.57 19.84 30.09 42.335.77 12.63 21.65 32.83 46.186.35 13.89 23.81 36.11 50.797.05 15.43 26.46 40.12 56.44961401E 2E 3E 4E 5E(1)407.94 17.36 29.76 45.14 63.492.3. Testing methodsThe as-built samples were evaluated to study the phase structures by the X-ray diffraction (XRD). A Bruker D8 X-ray diffractometer with a Cu Kα radiation (wavelength, λ = 1.5406 Å) was operated at 40 kV voltage with the current of 40 mA to identify phase structures.

A Bruker D8 X-ray diffractometer with a Cu Kα radiation (wavelength, λ = 1.5406 Å) was operated at 40 kV voltage with the current of 40 mA to identify phase structures. The radiation scanned sample surfaces continuously with a locked coupled step size of 0.01°, a speed of 1.17° min-1 and 2θ between 10° and 80°. The XRD profiles were then analyzed by Crystallographica Search-Match (CSM) program developed by Oxford Cryosystems [26]. To characterize morphology and microstructures, images on the surface of the blocks were taken using Carl Zeiss Sigma VP Scanning Electron Microscope (SEM). Average length, point distance and area of bright phases found on SEM images on nanoscale were calculated using ImageJ software developed by National Institute of Health (NIH). Energy Dispersive X-ray Spectroscopy (EDX) detector on the SEM model was used to identify the composition weight and atomic ratio of metal elements available in different samples.

Energy Dispersive X-ray Spectroscopy (EDX) detector on the SEM model was used to identify the composition weight and atomic ratio of metal elements available in different samples. No special preparation was needed for the testing of electrically conductive HEA samples and separation of the samples from the substrate was not required.3. Results and Discussion3.1. XRD AnalysisThe XRD measurements were carried out to examine the crystal structures of the SLM as- built CuCrFeNiTiAl HEA samples. Figure 3 shows the XRD patterns of the samples under various energy densities.Figure 3. XRD patterns of CuCrFeNiTiAl HEA samples under different energy densities. 3.2. Microstructure analysis962All XRD plots have been corrected for background noise and five strong peaks are visible in the 2θ range from 10° to 80°. The result shows that all three samples consist of five different crystal phases: hexagonal, orthorhombic, monoclinic, body-centered cubic (BCC), and rhombohedral.

The result shows that all three samples consist of five different crystal phases: hexagonal, orthorhombic, monoclinic, body-centered cubic (BCC), and rhombohedral. The alloy system comprises a major hexagonal solid solution matrix. Two major peaks are found at 2θ at 42.63°-42.85° and 43.52°-43.56° with rhombohedral and hexagonal crystal structures, respectively. High peaks responsible for rhombohedral structure in low, medium and high energy density areas, with intensities of 222 (2θ = 42.78°), 282 (2θ = 42.63°), 383 (2θ = 42.85°) respectively. Similarly, for hexagonal structure indicating high peaks in all three areas, respective intensities are 127 (2θ = 43.52°), 210 (2θ = 43.56°), and 404 (2θ = 43.54°). With increasing energy input during the laser melting process, the intensity of major XRD peaks and other minor peaks have increased.

With increasing energy input during the laser melting process, the intensity of major XRD peaks and other minor peaks have increased. However, the peak identified for BCC phase has higher intensity at mid-energy density area sample compared to high and low energy density area samples.Figure 4(a) shows an example of low magnification SEM pictures of the CuCrFeNiTiAl samples. It can be noticed that the surface morphology shows non-uniform microstructural distribution, which is considered due to the blending procedure. Additionally, though most of the high entropy alloy systems developed using casting process exhibit dendritic structural morphology, the CuCrFeNiTiAl samples in this study exhibited fine grained structures.Figure 4.

SEM pictures of CuCrFeNiTiAl sample of (a) 3C at low magnification; and samples of (b) 1A, (c) 3C, and (d) 5E at high magnification.963High magnification pictures of the samples were taken in specific regions of interest and multi-phase morphology was observed, as shown in Figure 4(b-d). The two distinct structures identified as bright area and dark area, with noticeable difference in shape, size and distribution. The amount of the bright phases shows increasing with an increase of energy density. In addition, it was observed that the bright phases in Sample 1A and 5E have polygonal shape while round- shaped bright phase was observed in sample 3C.Average length, point distance and area of bright phases are calculated using ImageJ software. The mean length of the bright phase is the largest in the high energy density area (Sample 5E) which is 186.24 nm and the smallest mean length is 122.92 nm which is found in the medium energy density area (Sample 3C).

The mean area of bright phases has increased with increasing laser energy input during melting from 8989.82 nm2 through 12857.71 nm2 to 13016.14 nm2. The point distances are measured to depict the distribution of bright phases throughout the microstructure. It is found that comparing to other two samples, the bright phases are positioned more closely in sample 5E with a mean point distance of 93.72 nm.3.3. EDX Element AnalysisEDX analysis was conducted for Samples 1A, 3C, and 5E to find out elemental compositions in both bright and dark areas upon different energy densities. Table 4 shows the composition of metal elements from EDX analysis given at weight percentage. Figure 5 shows the comparison of chemical composition in both bright and dark areas of the three samples with their nominal compositions.

Figure 5 shows the comparison of chemical composition in both bright and dark areas of the three samples with their nominal compositions. It is noticed that the EDX analyzed composition of each element was significantly different away from the nominal composition ratio, even though the elemental powder materials were pre-blended in equimolar ratio except for Al. Among all the principal elements, Cr has the lowest molar ratio and the highest molar ratio is observed in case of Ti. The molar ratio of Cr, Fe and Ni are lower than the nominal molar ratio (19 %), and Ti is about twice amount of the nominal values. However, Cu is the only principal element whose presence in different sample areas is consistent with the nominal molar ratio. In all samples, the molar ratio of Al is slightly greater than 5 %.Figure 5.

In all samples, the molar ratio of Al is slightly greater than 5 %.Figure 5. Chemical composition of (a) bright area and (b) dark area of CuCrFeNiTiAl samples.When comparing between the bright and dark areas, a significant difference is noticed in Sample 3C that the bright area of Sample 3C shows rich of Ti element, while Cu is rich in dark area. This elemental segregation can be explained by mixing enthalpy of principal elements. As964Cu has higher mixing enthalpy with other elements and mixing enthalpy of Ti is very low, Cu is repelled from the Ti-rich bright area [27].Moreover, though principal elements were mixed in equimolar ratio, the EDX analysis of the CuCrFeNiTiAl sample fabricated in SLM process found no equimolar ratio in any sample in this study. It can be indicated that insufficient energy was deposited into the powder blend.

It can be indicated that insufficient energy was deposited into the powder blend. The energy generated from laser during melting was from 5.29 KJ/mm3 to 63.49 KJ/mm3 and depending on the absorption rate of heat by the powder mixture, a portion of that energy is utilized for melting which might have not become sufficient enough heat to melt powder completely. Another reason was that the rapid cooling of the melt pool was not able to provide sufficient time window for the diffusion of the elements to form random solid solution.Table 4.

ConclusionThis study was intended to fabricate CuCrFeNiTiAl high entropy alloy by SLM process directly from the blend of six elemental powder materials. A sample matrix was produced on a 50.8 mm × 50.8 mm customized stainless-steel substrate, using different laser power and scan speed via full factorial experimental design. Three representative samples were selected for analysis and comparison in the as-built status. These three selected samples were then characterized using XRD, to study crystal structures, and using SEM and EDX to study the microstructures and elemental composition respectively. The conclusions drawn based on the investigation can be summarized as below:(1) Five crystal structures were observed in the SLM-fabricated CuCrFeNiTiAl alloy samples including: hexagonal, orthorhombic, monoclinic, BCC, and rhombohedral.

Thereinto, two major peaks appeared on the XRD graphs were found responsible for rhombohedral and hexagonal structures at angles of 42.63° and 43.56°, respectively.(2) Through SEM observation at a low magnification, the samples exhibited non-uniformmicrostructural distribution which might be caused during the blending process.(3) At a high magnification of SEM, the microstructures in three samples appear fine-grainedbright phases dispersed on dark phases.965(4) Shape, size and distribution of bright phases have a significant correlation with laser energy density under different printing parameters. While round-shaped bright phases in the sample using medium energy density had the smallest grain size, sizes of polygonal grains in the samples with low and high energy densities were comparatively higher.

The total area of bright phases increased with increasing energy input and the bright phases were most closely clustered in high energy density sample.(5) Though mixed in equimolar ratio, no equal compositions of the five principal elements were obtained based on EDX results, indicating incomplete diffusion during melting and solidification process.The multiple crystal structures observed in the samples strongly suggest that the heating process from the laser source was insufficient to create a time window that allows for full diffusion of the principal elements, and therefore, desired random solid solution was not obtained.

rect Slicing of STEP Based NURBS Models for Solid Freeform FabricationB. Starly, A. Lau, W. Sun Dept of Mechanical Engineering and Mechanics, Drexel University Philadelphia, PA 19104andW. Lau, T. Bradbury, A. Youssef, C. Gaylo Therics, Inc. 115 Campus Drive Princeton, NJ 08540 Reviewed, accepted August 13, 2003Abstract Direct slicing of CAD models to generate process planning instructions for solid freeform fabrication may overcome inherent disadvantages of using STL format in terms of the process accuracy, ease of file management, and incorporation of multiple materials. This paper will present the results of our development of a direct slicing algorithm for layered freeform fabrication. The direct slicing algorithm was based on a neutral, international standard (ISO 10303) STEP-formatted NURBS geometric representation and is intended to be independent of any commercial CAD software.

The direct slicing algorithm was based on a neutral, international standard (ISO 10303) STEP-formatted NURBS geometric representation and is intended to be independent of any commercial CAD software. The following aspects of the development effort will be presented: 1) Determination of optimal build direction based upon STEP-based NURBS models; 2) Adaptive subdivision of NURBS data for geometric refinement; and 3) Ray-casting slice generation into sets of raster patterns. Feasibility studies applying the direct slicing algorithm to example models and the generation of fabrication planning instructions involving multi-material structures will also be presented.Keywords: Direct slicing, STEP, NURBS, free form shapes, adaptive subdivision.1.

IntroductionLayered Manufacturing (LM) or Solid Freeform Fabrication (SFF), a new class of manufacturing techniques introduced during the mid 1980s have grown rapidly over the past decade because of its proven ability to reduce product development cycle time. In the current industrial Rapid Prototyping (RP) practice, the 3D CAD data are first converted to an intermediate STereoLithography (STL) format, a tessellation procedure where the model is approximated by triangles, sliced and then fabricated by the machine. But with the rapid growth of the RP industry, particularly RP towards rapid manufacturing, rapid tooling and biomedical applications, there has been a growing dissatisfaction with this format among the RP community due to the limitations inherent within the format. The geometric description used to represent solid CAD objects significantly affects the accuracy and quality of the final parts produced with this technology especially in the case of freeform shapes.

The limitation of the current STL format generated through a tessellation procedure of the CAD model can be summarized as follows [1]:162Tessellation involves approximation of surfaces with triangular facets which is undesirable in general, particularly when we are dealing with models that contain freeform shapes.As model precision demands become more stringent, the number of facets required to adequately approximate these freeform shapes will increase. This would account for huge STL file sizes, and thus increasing the chances of errors.Many CAD systems fail to generate valid model tessellations and often involve manual fixation of errors raising the need to improve tessellation procedures.

However, this calls for implementing robust and efficient procedures which may be difficult and would be computationally expensive to implement.STL format fails to include other design content within the model such as topology, internal material variations, multi-material regions.Manufacturers of free form surface models are in need of something other than the currently available STL format. Direct Slicing of CAD models without the intermediate STL format seem to be a promising approach in this direction. Direct slicing of the solid model keeps the geometric and topological robustness from the original data. Its advantages include greater model accuracy, pre-processing time reduction, elimination of checking and repairing routines, and file size reduction [2]. Since the direct mathematical formulation of the surface is used, the full data of the original solid modeler is therefore available and the loss that occurs during tessellation is avoided.

Since the direct mathematical formulation of the surface is used, the full data of the original solid modeler is therefore available and the loss that occurs during tessellation is avoided. Both direct slicing and adaptive direct slicing improve the accuracy and surface quality of the final RP parts. The approach also eliminates the verification and repairing processes, decreases human intervention, increases the robustness of data transfer, slicing and other preprocessing parameters. With this approach of direct input from computer aided design models, CAD and RP vendors would be able to sell their software/machines as a fully integrated CAD/CAM RP system.Accuracy was not an issue that was addressed upon in the early days of RP simply because the parts could only be used for prototyping and design verification purposes. A wealth of research is available in published literature [3-5] regarding STL slicing and the different methods through which it can be achieved.

A wealth of research is available in published literature [3-5] regarding STL slicing and the different methods through which it can be achieved. However, improving the accuracy of an RP part has become the focus of the RP community under the increasing need for prototyping functional parts with engineering properties and dimensional tolerances comparable to conventionally produced parts. One among the first to try out direct slicing was Rajagopalan et al [6], where they have directly sliced the Non-Uniform Rational B-spline Surfaces (NURBS) in an I-DEAS based CAD system. The process relied on I-DEAS to perform the slicing which made it package- specific. With regards to improving surface finish, variable thickness slicing methods (adaptive slicing) for handling peaks, flat areas and staircase effect have also been proposed by Dolenc and Makela [7].

With regards to improving surface finish, variable thickness slicing methods (adaptive slicing) for handling peaks, flat areas and staircase effect have also been proposed by Dolenc and Makela [7]. Jamieson and Hacker [2] developed a direct slicing algorithm based on the Unigraphics slice modules which directly sliced the model using first constant layer thickness. Consecutive contours were compared and if the difference was small, they were accepted. If not, a middle slice is created and the process of comparison performed again. The procedure is repeated until the difference between any two consecutive slice contours is either small or the minimum layer thickness has been reached. It has been recognized that in order to improve the surface quality of a part built by an LM technique, it is necessary to minimize the staircase effect which is inherent in all LM techniques.

It has been recognized that in order to improve the surface quality of a part built by an LM technique, it is necessary to minimize the staircase effect which is inherent in all LM techniques. In a more recent study, Weiyin et al [8] developed an adaptive direct slicing algorithm that operates directly on NURBS based models to generate skin contours and then uses a selective hatching strategy to reduce the build time of the model. Almost all of the papers published above in some way depend on external modeling packages to163perform the slicing which in turn limits the capability on the level of control and variety that can be achieved. A STEP-based transfer of model data to a process planning system has been made use of by Gilman et al [9] using commercially available software. A Boundary Representation (B-rep) solid model of a part is translated into STEP format for transfer of design data to the process planning software.

A Boundary Representation (B-rep) solid model of a part is translated into STEP format for transfer of design data to the process planning software. LM data is generated using faceted boundary representation and then transferred to the RP machine for prototyping. It has been recommended that the development of STEP specifically for the purpose of LM process planning due to its flexibility in terms of representing 3D CAD data as well as 2D contour slices thereby simplifying the standardization procedure with one common platform [1].The objective of this paper is to present a method on a direct slicing approach that is independent of any CAD modeling package and will make use of STEP as the starting input file of the model to be prototyped. We have currently focused on NURBS based freeform shapes to demonstrate the capability of the algorithm and the proposed methodology with regards to raster line pattern layout.

We have currently focused on NURBS based freeform shapes to demonstrate the capability of the algorithm and the proposed methodology with regards to raster line pattern layout. The following section will briefly detail our proposed methodology and some of the key steps in the process of direct slicing. This is followed by a review of the NURBS equations that need to be numerically solved to obtain the slice ray patterns during the slicing operation. The key steps of optimal orientation, adaptive refinement, root finding and evaluation of points are detailed. Section 5 gives several case studies example models that have been sliced to showcase the proposed method. The paper concludes with a summary and future research initiatives.2. Direct Slicing of NURBS using Ray CastingNon-Uniform Rational B-Spline (NURBS) are the industry standard tools for the representation and computer aided-design of freeform models [10] in the field of automotive design, ship design etc.

Central to the problem of slicing NURBS surfaces is the determination of intersection points between the slicing plane and the model. This is also a problem that is researched upon by the computer graphics community with regards to Ray-tracing of NURBS surfaces [11-14]. Ray tracing of free form surfaces determines the visible parts by constructing rays from the viewpoint through the pixels of the image plane into the scene. The rays are intersected with the surfaces of the scene and the first surface hit determines the color of the pixel. If the ray misses all objects, the corresponding pixel is shaded with the background color. Ray tracing handles shadows, multiple specular reflections, and texture mapping in a very easy straight-forward manner. This is important to the designer since it would help him assess the surface quality and texture and hence help in a better design process.

This is important to the designer since it would help him assess the surface quality and texture and hence help in a better design process. The basic approaches in most ray-tracing algorithms with regard to determination of the intersection points remain the same and only vary with regards to efficiency in terms of memory usage and the speed taken to ray trace a particular scene. Most ray tracing algorithms have performed while tessellating freeform shapes to gain speed and reduce complexity. However direct intersection using the exact mathematical equations has also been performed [14]. In our proposed methodology, we have used the same basic approach using the exact mathematical equations to deal with finding the intersection points which will be detailed in the following sections. We have however extended it to the slicing domain for use in the LM manufacturing scenario.

We have however extended it to the slicing domain for use in the LM manufacturing scenario. The difference lies in the fact that in LM, intersection points both in the form of entry and exit points need to be determined along with the vector layout pattern of the rays within the model. Rather than a tracing operation, a ray casting method is performed in order to perform the slicing procedure.164Within STEP files, solid and surface models may be represented as rational or non- uniform rational B-spline surfaces. Unlike STL files where the facet information of the triangle is used to obtain the slice contour, direct slicing works on using the exact mathematic representation of the freeform shapes in computing the slice contours or tool patterns.

A rational B-spline surface (10) is expressed parametrically in the form,+ 1+ 1nmÂÂvbubPW )( ik)(ijijjlvuS ),(=i= 1 + 1= 1 + 1jnmÂÂvbubW )( )(ijikjl(1)i= 1j= 1where parameters u and v range from zero to one, n and m the degree of the surface in u and v direction. The Pij terms are 3D net control points of the control polygon and Wij terms their corresponding weights, bik and bjl are B-spline basis functions of order k and l respectively.

The Pij terms are 3D net control points of the control polygon and Wij terms their corresponding weights, bik and bjl are B-spline basis functions of order k and l respectively. The B-spline basis functions are defined by the Cox-deBoor recursion formulas as given by:bj1 (s) = 1 if kuj ≤ u ≤ kuj+10 otherwisebl1 (s) = 1 if kvl ≤ v ≤ kvl+1 0 otherwiseandbjku )()( vb lk==u ((vkub ) - ( )1 )( kj - kujku -+ kj 1 ) kv b ( kl )( - l)1 kvkv+ kl 1lu )(j )( v++u )((kubu ) ( j )(1 - ku)1+++ 1 kj kuk++ 1 kj - ) bv -+ 1j)( vkv(++ 1 kl kv+ )(1( l kvk)1+ 1++ kl 1l(2)where the values of kuj and kvl are defined by the knot vector associated with the NURBS surface in the u and v direction respectively. The STEP file contains all information that is required to define the NURBS uniquely and a STEP reader may be employed to extract the relevant information.

The STEP file contains all information that is required to define the NURBS uniquely and a STEP reader may be employed to extract the relevant information. For further discussion on B-splines and their properties, the author refers them to [12].The core aspect of our approach is outlined in Figure 1. The NURBS surface is geometrically refined using an adaptive subdivision procedure to break them down into smaller domains of parametric values. Bounding boxes are then used to cover up the entire surface based on these smaller domains. The rays shoot out intersecting the model at several boxes. For a box that is hit, a root finding procedure is initiated to converge at the intersection point. The procedure is repeated for all rays that are cast onto the slice plane and for every slice plane that intersects the NURBS model.165Figure 1: Step by Step procedure on calculation of NURBS intersection points3.

Ray Casting of NURBS surfacesIn the ray casting approach, the bounding box of the model is first determined allowing the start position of the ray to be defined from any predefined corner of the box which then shoots out across intersecting the model. A ray is defined as having an origin ‘o’ and a unit direction vector ‘d’ and can be defined as:+=d*so Using the method followed by Kajiya [11], the ray r(s) can be rewritten as an intersection between 2 planes given by {p| P1.(p,1) = 0} and {p| P2.(p,1) = 0} where P1=(N1,d1) and P2=(N2,d2). The normal to the first plane is defined asr(s)N1=Ï ( Ô y Ì ÔÓ ,0( d,dzd x - d)0,,)yif |dx| > |dy| and |dx| > |dz| otherwiseN2 will always be perpendicular to the ray direction and the plane N1 , henceN2=dxN 1Since both planes contain the origin ‘ o ’, it can be deduced that P1.(o,1) = P2.(o,1) = 0. Thus,d1 = -d2 = -1N . o 2N.

Thus,d1 = -d2 = -1N . o 2N . oAn intersection point that needs to be calculated should satisfy the following two conditions,1P .(S(u,v),1) = 0 2P .(S(u,v),1) = 0The above equation needs to be solved using numerical means and we have employed the Bisection Iteration routine to determine the values of u and v that will satisfy (7). However before the root finding operation begins, a number of pre-processing steps are performed. In the steps that follow,166(3)(4)(5)(6)(7)we explain the details regarding optimal orientation, refinement using adaptive subdivision, generation of the bounding volumes, root finding, evaluation and identification of output points.Step 1: OrientationThe model orientation within the fabrication bed affects the build time, part strength and surface finish. Thus before the part is sliced, a minimization of certain objective criteria specified by the designer will be done to find the optimal orientation for slicing the model.

Thus before the part is sliced, a minimization of certain objective criteria specified by the designer will be done to find the optimal orientation for slicing the model. A number of orientation schemes have been devised. Some base their orientation with the largest convex hull of the object as the base [15], while some others orient the part based on certain critical features of the model [16]. We have used a scheme in which the model is incrementally oriented about user specified axes to obtain the least possible build height dimension. Given a set of n NURBS surfaces S(u,v) that is enclosed in a bounding box of height H (a function of orientation angle “_”), the objective function can be mathematically expressed as:Min Z = H(_) subject to 0 < _ < 360 (8)where Z is the build height of the model and _ is the orientation of the model with respect to the object coordinate axes. The optimally oriented NURBS faces then act as the input to the second phase.

The optimally oriented NURBS faces then act as the input to the second phase. The algorithm steps are as shown in Figure2.Step 2: RefinementFigure 2: Optimization AlgorithmRefinement or subdivision of the NURBS surface is the addition of more control points to a surface without changing its process. This process is implemented using the Oslo Algorithm [17]. The basic idea would be to take in the original set of knot vectors that make up the surface and add new knot values into them creating more number of control points corresponding to the new knot vectors. If the addition of the new knot values at the same parametric value where the number added is equal to the order of the curve, then the two new surfaces created will have the same shape as the original unrefined surface. They would eachFigure 3: Refined Process of NURBS using the Oslo Algorithm167have a set of control points at the region where they join. The procedure is illustrated in Fig 3.

They would eachFigure 3: Refined Process of NURBS using the Oslo Algorithm167have a set of control points at the region where they join. The procedure is illustrated in Fig 3. This is done for the following two reasons. Numerical methods work better and faster when the parametric domain is smaller and no multiple value roots exists within the patch. Secondly, by refining the mesh, the various sub-patches which are essentially NURBS by themselves can be enclosed using bounding volumes thereby enabling the slice algorithm to determine which sub-patch contains the actual root. This stems from the fact that it is easier to determine the collision of rays with primitive bounding volumes than NURBS surfaces.

This stems from the fact that it is easier to determine the collision of rays with primitive bounding volumes than NURBS surfaces. Hence by identifying the sub-patch in which the solution exists, the domain in which the numeric solver has to work is limited and hence results in better chances of finding the roots.The adaptive subdivision of the NURBS surface continues as long as a subdivision or flatness criteria is met. Regions that have more curvature are subdivided more than regions that are more or less flat and hence the name adaptive subdivision. Each new sub-patch contains all information that defines the NURBS and an appropriate ID is given to it. The refinement procedure is extensively used in the tessellation of parametric surfaces and has been studied extensively by a number of researchers.

The refinement procedure is extensively used in the tessellation of parametric surfaces and has been studied extensively by a number of researchers. However, our main criteria in refinement of the mesh opposed to the tessellation procedure is not to ensure accuracy in representation but more in guaranteeing that the convergence occurs within the refined sub-patch. In this regard, selection of the subdivision factor is an important step, and an appropriate value controls to a great extent the success of accurate slicing.Step 3: Boundary Volume Data structure GenerationAs the refinement procedure of NURBS progresses, the boundary volume data structure gets filled up; sub-patches are stored in the data structure along with a unique id. The main idea168Figure 4: General Slicing Algorithmbehind the storage of these sub-patches is the generation of the boundary volume. Boundary volumes are usually primitives that enclose the sub-patch completely.

Boundary volumes are usually primitives that enclose the sub-patch completely. Some of the candidate primitives that can be used are oriented boxes, spheres and parallelepipeds. The selection of a primitive depends on the tightness of fit and speed of intersection calculation with the ray. We have used in our approach axis aligned boxes that are oriented with the main coordinate axis of the model. Once they are axis aligned, the process of creating the boundary volumes is easier, and a ray-box intersection will involve less computation and hence speed up the process.The bounding box is created using the control net points created from the refinement stage of the process. An important aspect to note is that the control mesh of a NURBS patch will always enclose the surface and therefore any convex shape surrounding the control net would also enclose the surface. Figure 1 gives an illustration in 2D on how the boxes get generated.

Figure 1 gives an illustration in 2D on how the boxes get generated. Failure to enclose the sub-patch surface completely results in ray hits being missed and hence gaps in the sliced model.Step 4: Numerical Solution - Bisection Iteration RoutineThere are a variety of numerical methods that are available to solve for the intersection points. Among them include bisection algorithm, linear interpolation, Newton Iteration, and fixed point iteration [18]. Although the latter two methods are fast, in some cases they do result in solution divergence rather than convergence. Our problem reduces to finding the value (u*, v*) that corresponds to the intersection point of the ray and the NURBS surface. Though ray- tracing methods in model rendering employ Newton’s iteration since it proves to be a faster process, we have used the bisection iteration routine mainly due to its simplicity and robustness.

Time is not a critical factor in slicing as opposed to obtaining the roots of the equation.The surface sub-patch contained within the boxes that were hit by the ray are retrieved from the boundary volume data structure and passed onto the solver method. The routine works with iterating towards the solution of one variable. However, we need to solve for 2 variables (u, v) that satisfy the equation. This is achieved by keeping one variable constant and iterating towards the best value of the second variable that satisfies the equation. If the error in the points generated does not satisfy a tolerance, _, the first variable is incremented by a pre-defined amount and the procedure repeated. The process continues until either a solution is found or the limit of the first variable is achieved.We use two criteria to decide when to terminate the Bisection iteration routine.

The process continues until either a solution is found or the limit of the first variable is achieved.We use two criteria to decide when to terminate the Bisection iteration routine. The first condition is our success criteria: if we are closer to the desired point by some determined tolerance, _, given by:| F (u , v)| < _ (9)we then report a successful hit. The value of _ determines two aspects, first, the accuracy of our results and second, success in reporting a hit. A tight value might result in the routine reporting a miss and on the other hand a bigger tolerance will result in intersection point offset errors. The second condition is the failure criterion, meaning that if this condition is met, the routine exits reporting a miss. If during the routine, the error calculated is approximately same as the error in the previous iteration, the iteration exits reporting a miss provided the success criteria has not been met.

If during the routine, the error calculated is approximately same as the error in the previous iteration, the iteration exits reporting a miss provided the success criteria has not been met. This is mathematically written as169Fn(u,v) = Fn-1(u,v) subject to Fn(u,v) > _(10)where n is the nth step of iterationStep 5: Categorization of Intersection Points (Entry or Exit)Once the intersection points are found out for every ray that is cast onto the slice planes, these points need to be classified as an entry or an exit point. We have done this by calculating the normal of the surface at the point under consideration.

We have done this by calculating the normal of the surface at the point under consideration. The normal vector can be calculated by the cross product of the tangent vectors in the u and v direction evaluated at the point given respectively by equations (11) and (12) asTu =Tv =d ud d vd),( vuS),( vuS(11)(12)The normal vector at this point is then given by N = Tu x Tv(13)If it is assumed that the ray shoots across in the y direction, then a point is classified to be as: an entry point If Ny < 0; an exit point if Ny > 0; and an edge point if Ny=0. The points once classified are stored in a predefined format to be displayed on screen and for conversion to machine instructions.4. Implementation and Case Study examplesThe algorithms were implemented in C++ and tested for a variety of model shapes that included simple NURBS surfaces as well as complex curved ones.

Implementation and Case Study examplesThe algorithms were implemented in C++ and tested for a variety of model shapes that included simple NURBS surfaces as well as complex curved ones. Figure 4 gives the generalized algorithm for ray-casting of NURBS surfaces that includes all of the steps outlined above. Three of the test example cases are as shown below in Figures 6 to 8.Figure 5: Process Flow during FabricationAll the models were sliced at a constant layer thickness and ray offsets on the slice plane. Figure 5 details the data process planning steps right from the CAD file input stage to that of the final fabrication stage. For heterogeneous models, intersections of a ray to each geometric body are170calculated in sequence.

For heterogeneous models, intersections of a ray to each geometric body are170calculated in sequence. Thus the material associated with the geometric body being sliced can be captured and stored in the slicing data structure.Model 1: This model is relatively small with about 16 NURBS surface definitions and was extracted from the STEP input CAD file using a STEP reader. A subdivision factor of 0.01 and a tolerance value of 0.01 were used for adaptive refinement and root finding respectively. Figure 6(a)(a)(b)(c)Figure 6(a): Case Study Example Model 1 with 16 NURBS surfaces. Figure 6(b): Slice Raster Lines at a slice layer. Figure 6(c) : Entry and Exit Points displayed(a)(b)(c)Figure 7(a): Case Study Example Model 2 with 102 NURBS surfaces. Figure 7(b): Slice Raster Lines at a slice layer. Figure 7(c): Entry and Exit Points displayed.(a)(b)(c)Figure 8(a): Model 3 Layout – 135 NURBS surfaces.

Figure 7(b): Slice Raster Lines at a slice layer. Figure 7(c): Entry and Exit Points displayed.(a)(b)(c)Figure 8(a): Model 3 Layout – 135 NURBS surfaces. Figure 8(b): Sliced View layout Figure 8(c) : Sliced entry and exit points displayed171shows the model optimally oriented with the least height in the z-direction. Figure 6(b) shows a slice layout on the fabrication bed at the slice position defined by the slice plane as shown in Figure 6(a).Model 2: The second model contained 102 NURBS surface definitions, extracted through the same method using the reader and the data transferred to the algorithm. In this case a subdivision factor of 0.1 was sufficient for appropriate slicing with no perceivable slice errors during part layout. The same tolerance value of 0.01 was used for the root finding routine.

The same tolerance value of 0.01 was used for the root finding routine. Figure 7(a) shows the model optimally oriented and Figure 7(b) illustrating the sliced model part layout.Model 3: The third model is more complex not only in terms of having more number of NURBS surface but also in terms of its overall shape. This model depicts the capability of the algorithm to handle multiple exit, entry points as well as accurate slicing at edges of the model. A subdivision factor of 0.1 and a tolerance value of 0.01 was used for the process. Figure 8(a) illustrates the model and Figure 8(b) the corresponding slice layout on the fabrication bed. It can be seen that the model was sliced appropriately with key features detected at the slice plane shown.5. ConclusionDirect Slicing of CAD model promises to offset the disadvantages posed by the STL format. It does not involve any file repair routines and file sizes are much smaller to handle.

ConclusionDirect Slicing of CAD model promises to offset the disadvantages posed by the STL format. It does not involve any file repair routines and file sizes are much smaller to handle. It also facilitates the possibility of slicing multi-material volumes or heterogeneous models, a definite advantage over STL files. Since the exact representation is used, complete design information is carried over to the fabrication stage with no loss in information. Although direct slicing does take longer to slice, this problem can be offset by efficient algorithms in terms of reducing memory usage and faster computing power. A point to note is also a careful selection of the subdivision factor and the tolerance value. We have found that under continued trials, an intuition is developed on the selection of the right parameters which works best for the models.

We have found that under continued trials, an intuition is developed on the selection of the right parameters which works best for the models. Future research would involve the development of an adaptive direct slicing method on heterogeneous models and an appropriate slice layer format in terms of STEP that would be able to transfer slice information across different RP platforms. Effort would also be put into quantifying the exact accuracy obtained by direct slicing rather than using the STL format.6. References1. Anne L. Marsan, Vinod Kumar, Debasish Dutta, and Michael J. Pratt, “An Assessment of Data Requirements and Data Transfer Formats for Layered Manufacturing”, Technical Report NISTIR 6216, National Institute of Standards and Technology, Gaithersburg, MD. Jamieson, R., Hacker, H. "Direct slicing of CAD models for rapid prototyping" Rapid Prototyping Journal, Vol 1(2), 1995. 2.3. Luo R C.; and Y. W.

Direct SLS of powder compositions used for self-propagating high-temPerature synthesis. LV. Shishkovsky, A.G. Makarenko+, A.L.Petrov. P.N. Lebedev Physics IIl:stitute, Samara branch. , +Samara Technical University, SHS laboratory. Novo-Sadovaja st. 221,'443011 Samara, Russia. Fax: +7 (8462) 355600. E-mail: fian@transit.samara.ru.AbstractThe conditions of a direct selective laser sintering (SLS) were resea;rched experimental in a volume and by layers for the eutectic and near-eutectic powder compositions, which usually used for the self-propagating high-temperature synthesis (SHS) technology. It was shown a possibility to realise during SLS process a control exothermic reaction of combustion exactly in the focus ofew-Iaser beam on Nd-YAG. The necessary parameters of a laser influence (a power, scan velocity and diameter 'of the laser beam), a dispersion and composition of the powder mixtures were determined for the such regime.

The necessary parameters of a laser influence (a power, scan velocity and diameter 'of the laser beam), a dispersion and composition of the powder mixtures were determined for the such regime. The 3D parts from intermetallic compounds were created without any post-process procedures, that expands the functional opportunities of the sintering models. K~y words: direct SLS, eutectic powder compositions, self-propagating high-temperature synthesis (SHS), control exothermic reaction of combustion...IntroductionThe sele9tive laser sintering (SLS) of powder compositions is a world-wide spread method of creation of 3D parts in the SFF technology.

Initially developed for wax and plastic materials, today researchers are announcing about possibility directly to produce the volume parts from a polymer coating metal, bimetals and ceramic powder materials (see a pioneer work [I], Proceedings of SFF Symposiums and, at last, our works [2,3]).Better results are obtained when applying a liquid phase sintering process, based on a mix of two or more powder materials, with a high and low melting point, respectively. .During the part build, the phase with a low melt point fuses and binds a phase(s) with. a high melt point. But conversation of a such 'green part' to a real metallic one is demand the post-processing techniques. There are procedures of powder preparation, a sintering in the special conditions (control of a bed temperature and around atmosphere), a post processing annealing in an oven and/or infiltration gives in the result a relatively long and complex process.

It decreases the advantages of SLS method together with another its problems (a high porosity, a large shrinkage/distortion and etc.). No reject the all achievements the method of SLS at present time, we propose an alternative strategy of the research for the expansion of the uses opportunities of 3D parts and partial decision of above mention problems in the SLS process. It is a search of new powder compositions and a combination in a single technological process a some related processes. For example, the SLS can combine with a laser soldering [3], that allows to receive intermetallic compounds as in a paper [4] and for account of a high adhesion of the melt of a solder powder to non-melting phase to improve the properties of the volume models.In this paper, the approach by the combination of the SLS process and a self-propagated high temperature synthesis (SHS) is examined.

The investigated powder materials were a famous in SHS technology the mixtures on the basis Ni-AI and Ni-Ti in eutectic and near- eutectic compositions. A possibility of control exothermic reaction of combustion was91realised exactly in the focus of a laser beam during a selective laser scanning by the surface of those powder compositions.Materials and equipmentThe SLS process was conducted with use the scan laser beam from a computer controlled Nd-YAG. Laser worked in regimes with tum on (- 3 kHz) and tum off (i.e., cw) an internal resonator modulation. On the laser the replaceable lenses are used with focus distances f = 149 and 336 mm (a diameter of laser spots was d = 50 and 100 mkm, accordingly). A laser power changed in a range 1-16 Wand was supervised on the suiface of powder with help of a measurement device.

A laser power changed in a range 1-16 Wand was supervised on the suiface of powder with help of a measurement device. The control of laser beam with a computer had been used in a range of field 50x50 and 100xi00 mm (i.e., 1024xl024 pixels) and had allowed to conduct a scan by any predefind counter, reproducing thus by layer a 3D part. The duration of a laser scanning determined the times of laser beam transition from point to point, a calm deflector in each point and could be changed by software over a wide range of values.During SLS process the powders of the following marks are used (tab!. 1). The size of powder fraction was determined with a help of the sieve analysis and an optical microscopy on NIOPHOT-30.Table 1. Properties of used powder materials [2,5,6].Type of N 1 powder N.

A size of particles < 50 mkm was used. The following compositions were prepared for the realization of the next exothermic chemical reactions:3Ni + Al ~ NbAI + QI, Ni + 3Ti ~ Th Ni + Q2.(1) (2)It would occur to create the intermetallic compounds NbAI and TbNi. That systems are selected as the model systems, which usually are using in the SHS technology. For those test systems the basis characteristics are known: an ignition temperature - Ti, energy activation of chemical reaction - E, a heat effect of reaction QI,2 on the one mole of Ni (see table 2 and fig.l).92PhaseNi2Ah NiAl NhAI Melt (Ni+AI). Ti, (OK)9IO±25 925±16 -Table 2. E, (kJ/mole) - 135±27 115±25 48,24Q, (J/kg)I,9x106 9,7x106 2,4x106 1,3xl06DiscussionFrom the beginning, the powders are treated in a free pour volume, significantly more than a depth of a sintering mono-layer.

We had been researched the conditions of combined SLS and control SHS processes for an one laser passage depend of a laser power, beam diameter, scan time and powder composition. Under control regime of the SHS we understood a realisation of the exothermic reaction of a combustion without a heat explosion and only in the zone of laser influence (Le., in the so-called diffusion or kinetic regime of the SHS). 'It allows to determine the interval of sintering depths for the one laser passage under a minimal deformation on the layer.Figures 2,3 show the experimental results of the laser influence on the Ni-AI powder equilibrium composition. Under a small laser power P < 7 Wand a great velocity of laser beam v > 25 cm/s the sintering depths were a such little, that a sintered material looses from a touch. With an increasing of the power (fig.2) and decreasing of the beam velocity the depth of sintering stripe is growing.

With an increasing of the power (fig.2) and decreasing of the beam velocity the depth of sintering stripe is growing. In the interval of laser powers among -- 8,4-8,8 W (fig.2) and scan velocities -- 2,9-8,6 cm/s visually we had been seen a control exothermic reaction combustion exactly in the zone of laser passage. At last, under values P > 8,8 W , v = 25 cm/s (fig.2) and P = 8,8 W, v < 2,9 cm/s (fig.3) an exothermic chemical reaction had been gone over to the regime of explosion and a front of combustion began to spread in the all directions independently without a laser influence. It can consider those regimes a critical. In according of SHS theory, a diffusion regime of combustion, which characterises of a stationary distribution of a synthesis wave, transforms to the regime of a heat explosion. It needs to remark (fig. I), that on the heating stage at the T < T e = 640 0C only a Ni2Ah (1 phase) was arisen ~d diffusion grew on the contact surfaces of the Ni and Al particles.

It needs to remark (fig. I), that on the heating stage at the T < T e = 640 0C only a Ni2Ah (1 phase) was arisen ~d diffusion grew on the contact surfaces of the Ni and Al particles. Then at the range 1132 °C > T > Te, a'NiAI (s - phase) has been begun to create. When temperature is reached T =1638 °C and would occur eutectic melt, then on the cooling stage from this melt an intermetallic NiAl and NiJAI (s -phase) will crystallise. The laser influence parameters must be warm up so, that it would be sufficient a laser energy and a heat effect of exothermic reaction for the creation of those compounds.An increasing of size particles in the Ni-AI mixtures was shown, the SLS process took place without SHS. As a result, for the laser spot d = 100 mkm an optimal interval (P,v) was changed for the combined SLS and SHS processes.Also, we are investigated a near-eutectic Ni-AI powder compositions (mixtures - 1:1, 1:2, 1:4, 1:5).

Deviation from an equilibrium composition sharply narrows the optimal interval of process's combination and practically a laser sintering was seen. Sintering mono-layers were a very brittle.For the Ni-Ti powder composition (chemical reaction (2» the combination SLS and SHS was realised also on the regimes with v = 2,6 cm/s and P = 7,2 - 8,8 W. However, the sintering monolayers were brittle although they had depths compare with before researched93powder composition NI-Al. Addition in Ni-Ti composition a material with a low melt tenlperature (braze powder) decreased a brittleness.tu during which powder composition was under the laser influence and accomplished a transformation from the ignition regime to non-stationary a heat explosion regime. In the first case (fig. 1, P > 8,8 W) tu = d/v is equal 0,2 ms and in the second case (fig.2, P =8,8 W) tu = 1,7 ms.

of molecule collisions during a chemical reaction (1», c - a heat capacity, R - a gas constant (= 8, 3144 J/mole*K). A width of zone of exothermic chemical reaction inthe moment of the ignition is equal E RTi__1 * exp(--) J ,__1 * exp(--) J RT~__1 * exp(--) J E__1 * exp(--) J ,Iv Qpko[where Iv - a heat conductivity.It would be interesting to compare the times tu with a prediction of the theory. A calculation by equation (3) gives for the NiAI system tind = 41 ms and for the NhAl system tind =0,74 ms. Here we used the values from the Tables 1,2 and CNi = 462 Jlkg*K, CAl = - Iv (Ni3Al) = 20±2 W/m*K, Q*ko(NiAl) = 879 Jlkg*K [7]; (l4±2)xI016 Wlkg, Q*ko(NhAI) =(57±8)xI016 Wlkg [5]. It is seen, that the calculated periods of induction by the order are agreed with the tLI. Moreover, a tind for intermetallic compound NhAI is located in an experimental interval 0,2-1,7 ms. So we proposed, there is a great probability to receive NbAI phase.

Moreover, a tind for intermetallic compound NhAI is located in an experimental interval 0,2-1,7 ms. So we proposed, there is a great probability to receive NbAI phase. A calculation of the zone width (4) gives a next: Li (NiAI) = 430 mkm, Li (Ni3AI) = 82,3 mkm. It is also a very important result, because a laser sintering process is demand a great selectivity. The pointed out values shown, that for the NhAI phase a width of exothermic chemical reaction's zone is compared with a diameter of a laser spot d = 50 mkm and no more that two time greater than a size of powder particles. So we can consider, what a combination of SLS and SHS processes will exist successfully exactly by the counter of created 3D part. A time for period of induction for NiAI phase is not sufficient, then for the laser influence times a NiAI will not grow and a large Li (NiAl) (which is order of 8-9 the laser beam diameters) could not be taken to attention.

An X-rays analysis of treated powder compositions would be give a more exact result and we will plan to realise it.lv(NiAI) = 9±1 W/m*K,The 3D parts a simple form (a cub, spheroid and etc.) were created from the (Ni-Al) powder composition of equilibrium mixture without any post-processes procedures. Investigations of properties of those 3D parts will plan in future also, but now can say, that their durability is more high then a sintered parts.Conclusions 1. The combination of SLS and SHSprocesses in a single technological process was successfully realized for the eutectic and near-eutectic powder compositions on the basis of Ni-Al and Ni-Ti systems on the cw - Nd-YAG laser.942. Optimal regimes of the laser influence were determined for the support of the control exothermic reaction of combustion exactly in the focus of the laser beam. tLI with a theoretical period on induction is present 3.

tLI with a theoretical period on induction is present 3. A compares of laser influence times - for experimental determined laser parameters (P,v,d). It is shown, that those tLI are enough for synthesis of the intermetallic compound (NhAI phase) exactly in the laser effected zone. 4. A good selectivity could be achieved during combination of SLS and SHS processes. A width of control exothermic reaction's zone - Li (Ni3Al) = 82,3 mkm is compared with a laser beam diameter. 5. The 3D parts of simple form had been obtained, which contended an intermetallic compound system without any post-processes procedures.In a future research an attention must be given to investigate of intermetallic phases in the created 3D parts and those properties during the combination of SLS and SHS processes depend of laser influence parameters.References1. Bourell, D.L., Marcus, H.L., Barlow, J.W., Beaman, J.J.

gure 1: The LAMP process during a repair operation - The laser cladding head is shown on the right. A touch probe head is currently held in the machining spindle. A repaired H13 tool steel die component is held in the vise.requires additional shipping and processing beyond that of bulk steel, making the eﬀect more pronounced. The solution to this problem, particularly in the case of hybrid material addition and removal proesses, is to re-use the waste material from the removal portion of the process.

Microstructural characteristics are a major factor in determining the performance of a material. Solidiﬁcation microstructure is a function of the material composition and the temperature gradient in both space and time [1, 14]. For acceptable results, any attempt at using recycled materials will have to address contamination of machining waste.From a broad enough view, any product is composed of three basic components:2Energy - reﬁnement of materials, manufacturing, transportation • Materials - the stuﬀ the part is made of, expendables in the manufacturing process that do not become part of the productLabor - design, manufacturing, transportationRapid manufacturing addresses the labor costs through automation. American manufacturing relies on automation and outsourcing to keep labor costs down.

American manufacturing relies on automation and outsourcing to keep labor costs down. As energy costs rise, shipping costs rise, tipping cost scales back towards domestic manufacturing once more.Even if powders used in a laser based rapid manufacturing system are produced from secondary materials, they still will have a signiﬁcant cost of transportation from various entities in the supply chain (from the producer of the waste material, to a recycling facility, then to the powder manufacturer). Transportation costs are diﬀcult to generalize, however if reuse of on-site waste materials is possible, thus eliminating some transportation costs, a savings is clear. If the powders are atomized, then energy will have been spent melting the material in the production process as well.3 Using Recycled Materials in the LAMP SystemThe LAMP system at MS&T cannot directly feed machining chips without modiﬁcation to the existing systems.

For compatibility’s sake, a batch of Ti64 powder was produced from Ti64 ma- chining chips. A deposit using the recycled powder was compared to a deposit made with virgin Ti64 powder.3.1 Making PowderA batch of Ti64 chips milled from a laser deposited part by dry machining were used in producing a batch of recycled powder. A ball mill, seen in ﬁgure 3, from United Nuclear was used to crush the chips into smaller pieces. Ceramic media was used in an attempt to minimize contamination of the powder. The progress of the milling operation can be seen in ﬁgure 4. At the end of the 990 minutes alloted for the experimental run of the ball mill, less than 50% of the material in the mill ﬁt within the -70 +325 mesh spec needed for use in the LAMP system.

At the end of the 990 minutes alloted for the experimental run of the ball mill, less than 50% of the material in the mill ﬁt within the -70 +325 mesh spec needed for use in the LAMP system. Clearly a diﬀerent method of reﬁning the chips into powders is needed before this technique will be useful beyond the laboratory.Figure 3: United Nuclear 3lb ball mill used for reﬁning chips into powder.3(a) As-machined(b) 60 minutes(c) 90 minutes(d) 150 minutes(e) 210 minutes(f) 450 minutes(g) 720 minutes(h) 990 minutesFigure 4: Ti64 chips ground into powder with a ball mill. The progress of the milling was sampled at the intervals shown.3.2 Deposition ResultsThe deposition sample is shown below in ﬁgure 5. A diﬀerence between the two samples is im- mediately evident: The deposite made with the recycled powder is dull while the virgin powder deposit is shiny. The sample was sectioned with a wire EDM, mounted in bakelite, polished, and ﬁnally etched with Kroll’s reagent.

The sample was sectioned with a wire EDM, mounted in bakelite, polished, and ﬁnally etched with Kroll’s reagent. As shown in ﬁgure 6, this discoloration is not limited to the deposit’s surface. The exact nature of the contamination is unknown, but something has clearly been added to the deposited material. The list of suspects include the ceramic polishing media, the ball mill’s tumbler, and the AlTiN coating on the mill used to create the chips. An EDS analysis will be used to quantify the contaminants.4 Directly Feeding Machining WasteThe alternative to reﬁning chips into powders useable by the LAMP system is to enable the system to directly process machining waste. The LAMP team at MS&T is currently working to add multi- format (powder, chips, and wire) material capability to the LAMP system. This capability will allow for choice in source of the feed stock to be based on economics or simple availabilty rather than being limited to expensive powders.

This capability will allow for choice in source of the feed stock to be based on economics or simple availabilty rather than being limited to expensive powders. The ﬁrst new component in this initiative, shown in ﬁgure 7 is a prototype feeder capable of feeding either powder or chips.4Figure 5: Deposition of the recycled powder and original powder onto a Ti64 plate. Both samples are 4 layers with the following system parameters: 1.0kW laser power, 3 grams/minute powder feed rate, 15 inches/minute travel speed.(a) Deposition with virgin Ti64 pow- der.(b) Deposition with processed ma- chining chips.Figure 6: Deposition Results5 Concluding RemarksAlthough the ball mill is clearly not the method of choice for reﬁning of chips into powder, the LAMP team at MS&T is making progress toward reuse of its machining waste. An analysis of the contamination in the milled powder sample will determine the next course of action. The chip feeder will be completed in the fall of 2008.

An analysis of the contamination in the milled powder sample will determine the next course of action. The chip feeder will be completed in the fall of 2008. These additions to the LAMP facility will enable multi-source material usage, allowing the use of the most economic material for the application.6 AcknowledgmentsThis research was supported by the National Science Foundation grants DMI-9871185 and IIP- 0637796, and a grant from the U.S. Air Force Research Laboratory contract # FA8650-04-C-5704. The support from Boeing Phantom Works, Product Innovation and Engineering, LLC, Spartan Light Metal Products Inc, UMR Intelligent Systems Center, and UMR Manufacturing Engineering Program, is also greatly appreciated.5Figure 7: Proposed Design for a chip feeding system.References[1] William Hofmeister, Melissa Wert, John Smugeresky, Joel A. Philliber, Michelle Griﬃth, and Mark Ensz. Investigation of solidiﬁcation in the laser engineered net shaping (lenstm) process.

Liou. Automatic process planning of a multi-axis hybrid manufacturing system. In DETC, September 29 - October 2 2002.[10] Jianzhong Ruan and Frank W. Liou. Automatic toolpath generation for multi-axis surface In Proceedings of the 2003 ASME Designmachining in a hybrid manufacturing systemg. Automation Conference, Chicago, Illinois, September 2-6 2003. Paper No. DAC-48780.[11] Jianzhong Ruan, Jun Zhang, and Frank W. Liou. Support structures extraction for hybridlayered manufacturing. In DETC, 2001.[12] Todd Sparks, Vinay Kadekar, Gail Richards, Frank W. Liou, Venkat Allada, Ming Leu, Faisal Anam, and Siddharth Shinde. An advanced manufacturing workshop for high-school teachers and students. In Proceedings of the 2005 ASEE Annual Conference & Exposition, Portland, Oregon, June 12-15 2005.[13] Todd Sparks, Vinay Kadekar, Yogesh Thakar, Frank W. Liou, and Ashok K. Agarwal. Edu- cating high school students and teachers in rapid prototyping and manufacturing technologies.

Liou, and Ashok K. Agarwal. Edu- cating high school students and teachers in rapid prototyping and manufacturing technologies. In Proceedings of the 2004 ASEE Annual Conference & Exposition, June 20-23 2004.[14] Jos´e E. Spinelli, Ot´avio Fernandes Lima Rocha, and Amauri Garcia. The inﬂuence of melt convection on dendritic spacing of downward unsteady-state directionally solidiﬁed sn-pb al- loys. Materials Research, 9(1):51–57, 2006.7.

RECT-WRITE FABRICATION OF INTEGRATED, MULTILAYER CERAMIC COMPONENTSD. Dimos, P. Yang, TJ. Garino, M.V. Raymond, M.A. Rodriguez Sandia National Laboratories Albuquerque, NM 87185ABSTRACTThe need for advanced (electronic) ceramic components with smaller size, greater functionality, and enhanced reliability requires the ability to integrate electronic ceramics in complex 3-D architectures. For rapid prototyping and small-lot manufacturing, traditional tape casting and screen printing approaches are poorly suited. To address this need, we are developing a direct-write approach for fabricating highly integrated, multilayer components using a micropen to deposit slurries in precise patterns. With this technique, components can be constructed layer by layer, simplifying fabrication. It can also be used to produce structures combining several materials in a single layer. The parts are either cofired or sequentially fired, after each layer is deposited.

It can also be used to produce structures combining several materials in a single layer. The parts are either cofired or sequentially fired, after each layer is deposited. Since differential shrinkage can lead to defects in these multilayer structures, we are characterizing the sintering behavior of individua1layers. This technique has been used to fabricate devices such integrated RC filters, multilayer voltage transformers, and other passive components. The direct-write approach provides the ability to fabricate multifunctional, multimaterial integrated ceramic components (MMICCs) in an agile and rapid way.INTRODUCTIONHigh reliability passive components (Le., capacitors, resistors, and inductors) and packaging substrates are commonly made using multilayer ceramic constructions.

However, the desire to achieve further miniaturization, greater functionality, and enhanced reliability in advanced electronic and microelectromechanical systems is driving the development of more highly integrated components based on combining different functional ceramics together into complex 3-D architectures. Currently there are numerous examples achieved by electronic ceramic manufacturers, in which various materials have been combined to produce monolithic, multilayer ceramics with sophisticated functionality, such as filters [1,2] and solid-state dc-dc converters [3]. Similarly, significant work is being done to permit embedding of passive components into low-temperature cofireable ceramic (LTCC) packages.

Similarly, significant work is being done to permit embedding of passive components into low-temperature cofireable ceramic (LTCC) packages. This trend towards higher level integration, which is the passive component analog of an integrated circuit (IC), places increasing demand on fabrication processes and manufacturers.The usual technique for making these multilayer, multi-material, integrated ceramics is based on cofiring of laminates made from layers of green tapes with screen printed features, such as conductor traces. This manufacturing method has been refmed over decades of development, driven by the needs for low cost and high production volumes. While this approach works well for simple parts such as multilayer capacitors that contain only a single active ceramic, fabricating multifunctional components in complex geometries is a significantly more33challenging problem.

Consequently, commercial development of sophisticated integrated ceramics has emphasized a highly empirical approach, making the technology impractical for low-volume, specialty components.To overcome this limitation, we are developing a direct fabrication approach that simplifies the processing and provides greater flexibility than would otherwise be possible with tape casting and screen printing approaches. The goal is to provide a rapid prototyping and agile manufacturing approach to fabricating multifunctional, multimaterial integrated ceramic components (MMICCs). This work is based on the use ofa commercial micropen system [4,5,6] for depositing electronic-grade slurries in precise patterns. In this paper, we discuss several specific issues that are critical to the direct fabrication ofMMICCs by the micropen approach, such as the relationship between slurry rheology and print topology.

In this paper, we discuss several specific issues that are critical to the direct fabrication ofMMICCs by the micropen approach, such as the relationship between slurry rheology and print topology. Characterization of the sintering behavior· of thick films, which is critical for assessing cofireability, has also been carried out using two different approaches. Finally, a number of specific devices that have been fabricated using the micropen are illustrated.EXPERIMENTALMicropenDirect deposition of precise patterns is being accomplished using a commercial Micropensystem (Ohmcraft, Inc.), which is an automated printing device for ceramic slurries [5]. The micropen has a computer driven x-y stage that uses CAD file instructions, which permits easy on-line design changes in contrast to screen printing where a new screen is required for a pattern change.

It is also inherently capable of laying down multiple materials in a single layer, which cannot be done with conventional tape casting techniques. The micropen utilizes nozzles with an inner orifice diameter as small as 1 mil and an outer diameter down to 2 mil for high defInition patterns. A critical feature of the Micropen is the force feedback control on the pen tip, which is based on balancing the upward force on the pen due to the extruding slurry and the downward force applied by an electromagnet, as illustrated in Fig. 1. This control leads both to a very reproducible print topology and the ability to print over variations in the topography (i.e., height) of the workpiece. However, since the workpiece needs to support the force of the pen, any underlying slurry layer must be dried before printing on top of it.The other critical feature of the system is the pump block, which uses two internalchambers to provide smooth, continuous delivery of slurry.

The slurries are loaded into a syringe which screws into the pump block assembly. To get uniform and reproducible processing it is critical to eliminate air bubbles in the slurry, which is accomplished by centrifuging, and in the pump system, which is done by bleeding the pump block.MaterialsCommercial thick-film slurries were used to characterize the printing process and tofabricate components. The materials used in the work were: a crossover dielectric (Ferro 10 38N), a Ag conductor (Ferro 1039), and a Ru02-based resistor (Ferro 87-102). Slurry viscosities were characterized using a Brookfield viscometer with a cone and plate attachment.34QperationFeedback ControlLED and photodiode height detectorElectromagnet ~Cantilever spring pen tubeMagnet~Pen tip~1)2)foreet3)t force...Lower pen, record positionInitiate pumping to triggerSteady-state printing, constant forceFigure 1.

Schematic diagram of the micropen print head showing the photo detector system for height control and the electromagnet force balancing system. The lower drawings illustrate the steps to initiate slurry printing.SinteringTwo different systems were developed to characterize the sintering behavior of individualthick film systems. While detailed descriptions will be published elsewhere [7], schematic diagrams of the two systems are given in Fig. 2.,irrora).LamprrnSamPle·LUJ FurnaceIIiOltmeterTope b)Figure 2. a) Schematic diagram of system for measuring (lateral) sintering behavior of free standing films. b) Schematic diagram of system for measuring z-axis sintering behavior of laterally constrained films.35Figure 2a shows the arrangement used to determine the sintering behavior ofunconstrained, free-standing thick films.

Square samples were pattemed with the micropen onto teflon or mylar substrates, so that the thick-film sample could be removed from the substrate producing a free-standing thick film piece with an area of 4mm x 4mm and a thickness of 75-1 00 f.lm (2-3 print layers). The samples were fired on top of a quartz setter, which did not react with the samples during the measurements. The sintering behavior was characterized by continuously monitoring the x-y dimensions of the sample with the video recorder during heating.In contrast, Fig. 2b illustrates the arrangerpent for characterizing the sintering behavior offilms constrained in the x-y plane. These measurements also used samples with an area of 4mm x 4mm and a thickness of75-100 f.lm that were deposited onto alumina substrates.

These measurements also used samples with an area of 4mm x 4mm and a thickness of75-100 f.lm that were deposited onto alumina substrates. A dual laser beam system was used to determine the z-axis shrinkage of the film by comparing the deflection of a reference laser beam to the deflection of an incident parallel beam from a mirror positioned on the sample as shown in the inset of Fig. 2b. The laser beam deflections were continually monitored during heating using position-sensitive photodetectors. The optical path length from the sample mirror to the photodetector was -1.5 m, which enabled small changes in the angular deflection of the sample mirror and, thus, very small changes in the sample thickness to be accurately determined.

It should be noted that this system was developed after it was determined that a standard extensometer could not be used, since it applied enough pressure to the sample to locally deform it during firing.RESULTSSlurry viscosities as a function of shear rate are shown in Fig. 3, which demonstrate that all the thick film pastes are shear thinning. There is a wide range of viscosities for these materials at low shear rates. However, the range of viscosities converge at higher shear rates, which are typical of the printing process. The range of shear rates (dv/dr) typical of the printing process waso110100Shear Rate (l/sec)Figure 3. Viscosity vs shear rate behavior for the electronic thick films used.36estimated using the following equation for an incompressible, Newtonian fluid with no flow against the wall of the orifice:dv/dr = (4/n)(Q/R?) ,(1)where Q is the volumetric flow rate (typically - 10-5 cm3/sec) and R is the inner orifice diameter (2-10 mil).

The settling rates correspond to the high viscosity range.An attractive feature of the Micropen is that the print topology is fairly insensitive to theslurry viscosity. This feature is illustrated in Fig. 4, which shows profilometry traces obtained on line prints of the dielectric, conductor, and resistor using identical print conditions. However, the slurry viscosity at the low shear rates is critical for optimizing materials for either high definition patterns or for filled regions. The most viscous slurry (Ag conductor) leads to sharply defined traces, whereas the more fluid resistor paste flows together to give a relatively smooth surface for filled areas, which are typical of thick-film resistors.5001000150020002500X,flmFigure 4.

Profilometer traces showing that the print topologies for materials with different viscosities are very similar and good using identical print parameters.Figure 5 illustrates the sintering behavior, at a constant heating rate of20°C/hr, for thethree thick-film systems. The Ag conductor sinters at a much lower temperature than the other two materials, which is not ideal for cofiring compatibility. The free sintering rate is also higher than the constrained sintering rate, which is commonly observed and is due to internal stresses in the constrained system [8]. The shrinkage curves for the dielectric and resistor systems both show a rather abrupt onset of sintering (at 725°-800°C), which is typical for liquid-phase sintering systems above the liquidus. The temperature compatibility also promotes good cofireability.

The temperature compatibility also promotes good cofireability. In addition, the free and constrained shrinkage curves for these two materials are essentially identical, which is also consistent with a liquid-phase mechanism, since a liquid phase tends to limit stress buildup. It should also be noted that the total volume shrinkage is nearly equivalent for each material in both constrained and unconstrained sintering, which is anticipated and verifies good correlation between the two experimental methods.3750 - - Constrained~ 30-8, 20•Dielectric"""- .~ c: 10 'C .s::. CIJ01 0 .:-----:=-=-=-~-:-:-::--""-=-:':-=-----=-_::_::_-~ oTemperature eC)Figure 5. Measurements of the sintering behavior (% shrinkage vs temperature) at constant heating rate (20°CIhr) for three thick-film systems.

Measurements of the sintering behavior (% shrinkage vs temperature) at constant heating rate (20°CIhr) for three thick-film systems. Measurements were carried out with the two experimental setups to determine the constrained and unconstrained sintering behavior.Prototype ComponentsThe Micropen system has been used to prototype various electronic components. Figure 6a shows a R-C band reject filter fabricated with the micropen, based on the schematic circuit layout in Fig. 6b. The capacitors are parallel plate capacitors with a single (crossover) dielectric layer. The impedance response is shown in Fig. 6c, which indicates a 27 dB suppression at the reject frequency, which is given by fc = (nRC)/2. The sharpness of the attenuation is given by achieving good matched values for the capacitors and resistors.!]a)b)Out..-..00400] 300g Q) 200 § "0 8.100 e0 C)200100 ;g ~ ~ ..-.. 0..

0.. ('I) ()Q -100'-"27dB attenuation0~ t:-2001001000Frequency (kHz)tI.l ('I)Figure 6. a) Photograph of 4-layer band reject filter fabricated with the micropen, b) schematic diagram of band reject filter, c) impedance response for a typical micropen band reject filter.The capability of the Micropen to accommodate various topographies has also been utilized to build a multi-turn voltage transformer structure, as shown in Fig. 7a. The transformer is built with an outer winding of six turns and an inner winding of three turns so that the device can be used for 2: 1 and 1:2 voltage conversion. This construction relies on the ability of the micropen to38print at a range of heights and would be extremely difficult to fabricate using standard tape/screen printing methods. This design works well, but the conversion efficiency (Fig.

This design works well, but the conversion efficiency (Fig. 7b) of the initial prototype is relatively poor since a low permeability dielectric was used for the prototyping, rather than a high permeability ferrite. In addition, we have successfully fabricated a flat eight turn solenoid (L -- 1 f.!H), and a multitap voltage divider, which are shown in Figs 8a and 8b, respectively. While these devices are all printed on alumina substrates, the process has also been used to fabricate components on tapes and on mylar and teflon substrates that allow free standing components to be built.a)til(6mm.-0.02 >: >- 2 J~" A:H f.~. ~ -.- v, 8' ~ E0 • tJ V \.1 \J 0.00 ~ V V V V ,& ~ -0.02 0~.r,,1 MHz1 \ "f't,[ ~ I-t -2o12 Time,J.S34b)(l.)Figure 7. A) Picture of multiturn voltage transformer fabricated with independent Ag spiral windings (6 outer turns, 3 inner turns) and crossover dielectric. Plot of input vs output voltage for the transformer.5mm ..

Plot of input vs output voltage for the transformer.5mm .. ,.4mma)b)Figure 8. a) Flat solenoid fabricated by connecting lower half windings and upper half windings around thick-film dielectric (L(@ 1 MHz) -- 1 f.!H), b) Multitap voltage divider for 10:1, 7:1,4:1, and 2: 1 voltage division. A 10kQfsq resistor with Ag electrodes is used.CONCLUSIONSWe have demonstrated a direct-write approach for fabricating highly integrated, multilayercomponents using a micropen to deposit slurries in precise patterns. With this technique, components are constructed layer by layer, simplifying fabrication. The quality of print features depends on slurry viscosity; however, the micropen can use a wide range of slurry viscosities39with minimal optimization. Since differential shrinkage can lead to defects in these multilayer structures, we have developed two techniques for characterizing the sintering behavior of individual layers, which should help in assessing cofireability.

Finally, the micropen technique has been used to fabricate devices such integrated RC filters, multilayer voltage transformers, and other passive components. The direct-write approach provides the ability to fabricate multifunctional,multimaterial integrated ceramic components (MMICCs) in an agile way with rapid turnaround.ACKNOWLEDGMENTSThe authors would like to acknowledge excellent technical assistance from I. Alderete.This work was supported by the United States Department of Energy under Contract DE-AC04 94AL85000. Sandia is a multiprogram laboratory operated by the Sandia Corporation, a Lockheed Martin Company, for the United States Department of Energy.REFERENCES 1. H. Hayashi, T. Ikeda, and S. Nishigaki, ISHM '91 Proceedings, 508-512 (1991). 2. H. Mandai, K. Wakino, H. Okamura, and 1.P. Canner, Ceram. Trans., 15 (Mater. ProcessesMicroelectron. Syst.), 391-404 (1990).3. T. Minoru, Y. Mochizuki, A. Nakano, and H. Kobnke, ISHM '92 Proceedings, 269-274 (1992).4.

rected Energy Deposition Processing-Performance Relationship of AF9628M. Yang1, M. M. Parvez1, T, Sparks3, S. Babalola2, J. W. Newkirk2, K. Chandrashekhara1, L. N. Bartlett2, and F. Liou11Department of Mechanical and Aerospace Engineering 2Department of Materials Science and Engineering Missouri University of Science and Technology, Rolla, MO 65409 3Product Innovation and Engineering, LLC, St. James, MO 65559AbstractAF9628 low alloy steel is a novel steel known for its low cost, high hardness, and outstanding tensile performance. However, the processing methods of AF9628 have been less studied in the additive manufacturing field. As the balance of hardness and tensile results is closely related to the cooling process during manufacturing process window and its relationship to resultant tensile properties was explored. By using the DED method, specimens of this steel were successfully fabricated, and tensile test results were obtained.

By using the DED method, specimens of this steel were successfully fabricated, and tensile test results were obtained. The AF9628 steel can be manufactured for tailored properties with the DED process by controlling the cooling method. This work was funded by ARL - GVSC under cooperative agreement W911NF-20-2-0251.Keywords: Directed Energy Deposition (DED), AF9628 Low Alloy Steel, Laser Power, Scan speed, Cooling Rate, Micro-tensile Test (YS, UTS, Elongation).IntroductionAF9628 is a novel alloy invented by the U.S. Air Force. Its high tensile strength and low cost enables it to be an outstanding material. The AF9628 has been widely studied with selective laser melting (SLM) process. A mean YS (yield strength) of over 1500MPa, a mean UTS (Ultimate tensile strength) of over 1700Mpa, and a mean elongation of over 10% were tensile tested along and perpendicular to the build direction [1].

However, the AF9628 low alloy steel has been rarely studied in the DED (direct energy deposition) process. It is likely that the tensile strength (YS, UTS, elongation) of the AF9628 alloy steel can be further improved. Once it is improved, the AF9628 low alloy steel can also become an ideal material for 3D rapid repair. This paper includes the study of the tensile results from the AF9628 DED process.Experimental SetupThe AF9628 low alloy is developed by Dr. Rachel Abrahams. This type of novel alloyinterests the industry for its ultra high-strength and low cost. As costly element such as Ni is limited, the cost of AF9628 is lower than the conventional alloy steel. It is also highly weldable since it does not contain any high melting point elements such as W or Ti. All these properties make the AF928 steel an attractive material for additive manufacturing process, which makes it doable to produce parts of more complex geometry.

All these properties make the AF928 steel an attractive material for additive manufacturing process, which makes it doable to produce parts of more complex geometry. [1]450Material:The AF9628 low alloy powder used in this study was the commercial AM powderproduced by Powder Alloy Corporation. The AF9628 particle size was analyzed by the AFA mode on the ASPEX Machine. The powder images (shown in Figure 1a.) were taken ranging from x120 through x 200. As shown in these images, the AF9628 powders purchased for this study were not perfectly spherical. [12,14] Some of the particles were in an elliptical shape. AF9628 Particle Size Distribution:The particle distribution is shown in Figure 1b. There were 5028 particles analyzed insize. According to Davg, about 39% of the particles were very fine(<10µm) 45% of the particles were fine(10-20µm)13 % of the particles were 20-60µm, and ~3% of the particles were >60µm. The particle size majorly ranges between 5-20 microns.

The particle size majorly ranges between 5-20 microns. The max population was found between 5-10 microns.500 microns500 micronsFigure 1a. AF9628C powder SEM image at higher and lower magnificationsFigure 1b. AF9628C Particle Distribution451The AF9628 low alloy powder used in this study was the commercial AM powderproduced by Powder Alloy Corporation. The chemical composition from literature review is listed below for reference (Table 1).Table 1 Chemical composition by weight percent of AF9628 [2]Mn Element Weight (%) 0.028 0.28 <0.005 0.0002 0.7 Al Element Ni 0.009 0.004 <0.035 Weight (%) 1.1OCSHSi 0.9 TiCr P <0.01 2.8 P+Sn+As+Sb FeMo Cu 1V 0.12<0.1BalanceThe Powder density tester used in this study was “Ultrapyc 1200e” the results are shownin Figure 1c.Figure 1c. Ultrapyc 1200e. Experimental settings: Nitrogen gas, Pressure: 10 PSI, Pulse: 3, Run: 3 Randomization applied: 21.257 gm, 12.663 gm, 17.668 gm. 𝐺𝑎𝑠 𝑝𝑜𝑟𝑒 𝑑𝑒𝑛𝑠𝑖𝑡𝑦: 0.475 ±0.119 (%).

Ultrapyc 1200e. Experimental settings: Nitrogen gas, Pressure: 10 PSI, Pulse: 3, Run: 3 Randomization applied: 21.257 gm, 12.663 gm, 17.668 gm. 𝐺𝑎𝑠 𝑝𝑜𝑟𝑒 𝑑𝑒𝑛𝑠𝑖𝑡𝑦: 0.475 ±0.119 (%). Theatrically powder density: 7.818 g/ccExperiment procedure and design:The AM system used in this study equipped a maximum 1000w fiber laser with a 2mmdiameter laser beam, an electrostatic wheel feeder with the feed rate set at 1.8 g/min. All the equipment was controlled by the LinuxCNC software.This study was focused on single track geometry of dimensions 45 mm x 2.5 mm x 15mm. The layer thickness was 0.2 mm (75 layers). The simply left and right moving tool path shown in Figure 2 was used in this study. To control the delay time and keep the gantry moving at a constant jogging speed, the gantry was programmed to move for a pre-calculated extra distance of 17 mm, so that the gantry jogs left and right at the scan speed of 200 mm/min to keep the 10 seconds delay time.

Once the gantry jogs to the very end of each direction the gantry lowers for 0.2 mm (set up thickness) then goes back to either the left or right edge of the geometry end. Once the gantry arrives at the edge of the geometry, the deposition of the next layer starts.452Figure 2. Single-track tool path. Delay calculation: Scan speed = 200 mm/min delay time = (17x2) / (200/60) = 10 secondsThe substrate used for AF9628 single track deposition was SS304 of a dimension of 5mmx 10mm x 55mm. A 350w laser power was used for preheating and the experiment parameters are shown in Table 2.Table 2.

A 350w laser power was used for preheating and the experiment parameters are shown in Table 2.Table 2. Single track DED deposition process parameters of samples #1 through #11Sample #1 2 3 4 5 6 7 8 9 10 11Layer width Layer Thickness Scan Speed Laser power Delay(mm)2.5(mm)0.2(mm/min) 200 200 150 200 250 150 200 250 150 200 250(W) 350 400 300 300 300 350 350 350 400 400 400(s) 10 10 no delay no delay no delay no delay no delay no delay no delay no delay no delayTo avoid the sample being oxidized, argon supply was kept on for at least 15 mins tomake sure the sample cools down below the oxidization temperature before collecting the sample.453Characterization:After the micro tensile samples were tested, the broken pieces were collected and mounted with epoxy for micro-structure observation. The polisher used in this study was MEPREP3 PH-3. The samples were polished to 0.05.

The polisher used in this study was MEPREP3 PH-3. The samples were polished to 0.05. After polishing the samples, the SEM images were taken under with EBSD to investigate the phases of sample 1 and sample 2.Mechanical Testing:Once the sample was deposited, the single-track deposition as well as the substrate were sliced vertically with EDM into 1 mm thickness pieces, the MT2 samples were extracted along the build direction side by side shown in figure 3(a). Rough finish (polish) with 800 grit sandpaper to remove the surface oxide, take the measurement with the micrometer for the width and thickness of the middle flask of the MT2 samples, and record the dimensions of the micro- tensile sample. The theorical dimensions of the MT2 samples are shown in figure 3(a).

The theorical dimensions of the MT2 samples are shown in figure 3(a). Repeating the process for all pieces, once the dimension table of the MT2 samples was made, then the samples were prepared.Sliced AF9628 deposition after MT2 samples extractionSliced SS304L substrate(a)(b)Figure 3(a). Locations where the micro-tensile samples were extracted from a 1mm thick slice Figure 3(b). Dimensions of the micro-tensile samplesThe tensile tester used in this study was INSTRON accompanied with an ex-tensile meter and Bluehills3 control program. It has a 50 kN maximum stress. The tensile test parameters were 0.015mm/mm/min test speed and 1.5mm/min rate 2.After the micro-tensile samples were tested, the debris were collected (shown in Figure 4a) and prepared for Vickers-hardness test. The hardness tester used in this study was “Struers Duramin5”.

The hardness tester used in this study was “Struers Duramin5”. On each piece of debris, three locations (shown in Figure 4b) were put along x-axis, perpendicular to the build direction on the wide end of each piece, three indents were put at each location to calculate the mean local hardness value. In total, twelve locations (36 indents) of each sample were tested for the global hardness value. All indents were tested by the parameters set at 9.81N 10 seconds and 40x magnification shown in table 6.454(a)(b)Figure 4a. Four pieces of prepared micro-tensile debris of sample 2 in one epoxy mount for hardness test Figure 4b. Three Vickers-hardness tested locations on one piece of micro-tensile debris labeled in blue color.Results and DiscussionMicrostructure:The porosity morphology shown in figure 5a sample 1 could indicate lack of fusion defects as the irregular defects were observed.

In figure 5a sample 2 the porosity is evenly distributed, which could indicate outgassing in this higher energy density specimen.As what is shown in figure 5b, the phase mappings of sample 1 and 2 are showingdifferent phase morphologies. Comparing with sample 2 the phases size of sample 1 are more uniform.Build DirectionBuild DirectionIrregular DefectsSample 1Sample 2Figure 5a. Sample 1 & 2 EBSD fore scattered detector (FSD) images for porosity comparison455Sample 1 phase colorSample 1 IPF colorSample 2 phase color Sample 2 IPF colorFigure 5b. Sample 1 & 2 phase mappings for phase size comparisonIn this study, ferrite and cementite phases were specifically focused to find Bainite. Aswhat is shown in Table 3, assuming the total carbon content of both samples is similar, the lower rate of a discrete carbide phase in sample 2 could indicate a higher prevalence of bainite than in sample 1.

Potentially more austenite phases are transformed into Bainite during the transforming stage of the cooling process in sample #2.456Table 3. DED sample 1 and 2 mapping analysisSample 1Sample 2Phase Name Fe-BCC Fe-FCC Fe3C Zero Solutions *Grain Diameter Median [µm]Phase Fraction (%) Phase Count Phase Fraction (%) Phase Count49.67 0.25 4.53 45.5597025 497 8846 8897159.18 0.29 2.95 37.5999581 485 4959 632451.50.9Mean grain diameter analyzed with phases over 10 pixelsSample 2's highest frequency bin (shown in Histogram 1) is the sub-micron category, indicating that the median grain size is much smaller than sample 1. The smaller median grain size in sample #2 supports the tensile result of sample #2 tested a higher UTS than sample #1 [13].(a) (b)Histogram 1. Sample #1 & #2 Grain Diameter Distributions [µm]Tensile PropertiesEleven samples were tested under the same tensile test parameters in this study.

Sample #1 & #2 Grain Diameter Distributions [µm]Tensile PropertiesEleven samples were tested under the same tensile test parameters in this study. Bothsample 1 and sample 2 were deposited at 200 mm/min and 10 seconds delay between each layer, but sample one was processed with 350 W laser power, and sample two was processed with 400 W laser power. The tensile results of sample 1 and sample 2 were shown in Figures 6a, 6b and compared with other tensile results in table 5.457Sample 1 and sample 2 were both tested for over 900 MPa yield strength and a relativelyhigh elongation of over 20% [14,15] as shown in Figure 6a and 6b. Sample one was tested slightly higher than sample 1 based on the calculated mean YS values and mean elongation percentages.

Sample one was tested slightly higher than sample 1 based on the calculated mean YS values and mean elongation percentages. However, a mean ultimate-tensile strength of over 1800 MPa was tested in sample two, which was about 400 MPa (28%) higher than sample 2.) a P M(s s e r t S e l i s n e T1500 1250 1000 750500 250 00.00Specimen 1Specimen 2Specimen 3Specimen 4Specimen 5Specimen 60.10 Tensile strain (mm/mm)0.050.150.200.25Scan speed Layer width Layer thickness Laser Power200 mm/min 2.50 mm 0.20 mm 350 WUTS YS (MPa) (MPa) Mean 975 1365 Median 972 1349 35 106S.D.Elongation (%) 22.8 22.9 1.90Figure 6a.

None of the nine consciously deposited samples was tensile tested for an over 1600 MPa UTS or an over 15% elongation. Sample #5 and sample #11 tested a mean UTS value close to 1500 MPa. Relatively uniform elongation percentages were tested on the no delay samples. The elongation range was between 6% to 8%.

Sample #5 and sample #11 tested a mean UTS value close to 1500 MPa. Relatively uniform elongation percentages were tested on the no delay samples. The elongation range was between 6% to 8%. The yield strength distribution was tested around 1000 MPa, and the highest mean yield strength was tested 1100 MPa at sample #5.458300 W Laser Power350 W Laser Power) a P M(h t g n e r t se l i s n e T1,5001,0005000100200 Scan speed (mm/min)150250300151050)%(n o i t a g n o l E) a P M(h t g n e r t se l i s n e T1,5001,0005000100250 200 150 Scan speed (mm/min)300151050)%(n o i t a g n o l EUTSYSElongation (%)UTSYSElongation (%)(i)(ii)400 W Laser Power) a P M(h t g n e r t s1,5001,00050015105)%(n o i t a g n o l Ee l i s n e T0100250 200 150 Scan speed (mm/min)3000UTSYSElongation (%)(iii)Figure 6c.

Tensile results of (i) 300W, (ii) 350W, (iii) 400W laser power no delay samples at different Scan SpeedsSample one and sample two tested relatively the same yield strength as the cast AF9628steel shown in Figure 6d, and the UTS of sample 1 was tested relatively the same as the cast AF9628. However, the UTS of sample 2 was 400 MPa higher than the UTS of the cast AF9628, and both sample 1 and sample 2 were tested about 7% higher than the cast AF9628 in mean elongation.1400) a P M(s s e r t se l i s n e T12001000800600400200Specimen 1 Specimen 3 Specimen 5Specimen 2 Specimen 4UTS (MPa) 978 1339 Mean Median 983 1356 41 150 S.D.YS (MPa)Elongation15.4 15.0 7.10000.050.10.150.20.25Tensile strain (mm/mm)Figure 6d.

Cast AF9628 stress-strain curve and tensile results459Sample one deposited in this study was tested about 100 MPA lower in YS comparedwith the SLM processed AF9628 tested at both vertical direction and horizontal directions. However, the UTS of the SLM processed AF9628 was about 100 MPa lower than sample 1 and 500 MPa lower than sample 2., and the mean elongation of sample one and sample two were both 50% high than the SLM processed AF9628 compared with [5] shown in Figure 6e.The YS in [1] shown in table 4. was about 500 MPa higher than either sample 1 or sample 2, and the UTS of [1] was 300 MPa higher than sample 1 but 100 MPa lower than sample 2 in all tested directions. However, the elongation of sample 1 and sample 2 were both about twice of the elongation percentage of the [1] results in all directions.Figure 6e. AF9628 SLM process stress-strain curve from [5]Table 4.

AF9628 SLM process stress-strain curve from [5]Table 4. AF9628 SLM process tensile results from [1] where TD-SD refers to XY plane (horizontal direction) SD-BD refers to YZ plane (vertical direction) BD-TD refers to XZ plane (vertical direction)Table 5. AF9628 tensile comparison of different processes in energy density, ultra-tensile strength, yield strength, and elongationAF9628 Processes DED (No Delay) DED (10 S Delay) CastED (J/mm3)240240n/aUTS (MPa)130918061339YS (MPa)988958978Elongation (%)6.321.815.4SLM [5]Not given1700153011.5SLM [1]60.1142011007.71460The ductile fracture surfaces shown in Figures 7a (i) and 7a (ii), and Figures 7b (i) and 7b (ii) were observed at both sample 1 and sample 2. No sign of brittle surface was observed at sample 1 or sample 2. Sample 1 and sample 2 have a similar surface texture at x500 magnification.

No sign of brittle surface was observed at sample 1 or sample 2. Sample 1 and sample 2 have a similar surface texture at x500 magnification. The overview at low magnification of sample 1 fracture surface was observed to be finer and slightly more uniform than the overview of sample 2 fracture surface. However, a larger pore defect was seen at the corner of sample 1 fracture surface. Sample 1 had a higher defect density than sample 2, which resulted in sample 2 being tested with a much higher UTS than sample 1 in the tensile test.Pore Defect(i)(ii)Figure 7a (i). Sample 1 fracture surface SEM image overview Figure 7a (ii). Sample 1 fracture surface SEM image at the conner of the fracture surface(i)(ii)Figure 7b (i). Sample 2 fracture surface SEM Image at x110 magnification overview Figure 7b (ii).

Sample 1 fracture surface SEM image at the conner of the fracture surface(i)(ii)Figure 7b (i). Sample 2 fracture surface SEM Image at x110 magnification overview Figure 7b (ii). Sample 2 fracture surface SEM Image at x500 magnification at the conner of the fracture surface461Hardness PropertiesSample 1Sample 2Sample 5Sample 8Laser Power (W)350400300350Scan Speed (mm/min)200200250250Delay B/W Layers (s)10No DelayPress Force (N)9.81Press Time (s)10Magnification40xMean Hardness (HV)437429480481Global S.D.88109Table 6. Samples 1 2 5 8 mean Vickers-hardness values referring to process parameters and test parametersAll the 11 samples were tested at room temperature for Vickers-hardness, and all samples were tested for relatively high Vickers-hardness values. [6,7,8,9,10,11]. The results of samples 1 2 5 and 8 were selected for comparison, their hardness results are shown in table 6.

[6,7,8,9,10,11]. The results of samples 1 2 5 and 8 were selected for comparison, their hardness results are shown in table 6. Sample 1 and sample 2 were deposited with 10 seconds delay between each layer, and sample 5 and sample 8 were without delay between each layer. Based on the global mean hardness value, sample 1 was closer to sample 2 in hardness, both were around 430 HV. Sample 1 was tested 8 HV higher than sample 2. Both sample 5 and sample 8 tested about 480 HV, and both were 35 HV higher than sample 1 and sample 2.The parameter differences of laser power and scan speed played a minor factor in thevariance of sample Vickers-hardness. The factor leading to the major difference in Vickers- hardness results can be the different cooling rates. When the sample was deposited with 10 seconds delay, the sample cooled down more compared to a layer-by-layer process.

When the sample was deposited with 10 seconds delay, the sample cooled down more compared to a layer-by-layer process. The temperature was further decreased, the grains can be formed more completely when each layer was scanned comparing with the samples deposited without delay between each layer. The cooling rate of the samples deposited without delay was more compared to the cast process, as the cooling time between each layer was shorter, the whole part cooled down as an entire body, hence resulted the sample in a higher Vickers-hardness.462ConclusionsThe AF9628C low steel alloy tensile property was related to its hardness property. Onerelatively increases as the other relatively decreases. The cooling rate was a significant factor for AF9628C low steel alloy to weigh the properties. A process of 10 s delay between each layer resulted in higher values in UTS and elongation, as well as lower values in YS and hardness.

A process of 10 s delay between each layer resulted in higher values in UTS and elongation, as well as lower values in YS and hardness. A process of no delay between each layer resulted in higher values in YS and hardness, along with lower values in UTS and elongation.Tensile PropertiesBoth sample 1 and sample 2 with 10 seconds delay were tested over 20% elongation, whereas the samples processed with no delay were all tested less than 10% elongation. An over 1800 MPa mean UTS was tested at Sample 2. However, the YS of sample 1 and sample 2 were both tested about 100 to 150 MPa lower than other samples processed with no delay between each layer. As the longer delay between each layer, the more heat will be emitted and conducted, the alloy grain can be more evenly formed.

As the longer delay between each layer, the more heat will be emitted and conducted, the alloy grain can be more evenly formed. As the alloy grains were formed evenly, the stress was more uniformly distributed on the sample body and less likely to accumulate over the break threshold at one point as the elongated distance increased, hence higher elongation percentage and UTS were tested at sample 2.Hardness PropertiesSample 1 and sample 2 were tested relatively lower Vickers-hardness values thanthe samples processed with no delay between each layer. The difference in Vickers- hardness results can be sourced to the different cooling rate. The over low melt pool cooling rate during solid-state phase transformation transformed mostly Austenite into Pearlite and the over high cooling rate during solid state phase transformation transformed Austenite mostly into Martensite. Both the Pearlite and Martensite can result the sample in higher hardness results [16,17].

Both the Pearlite and Martensite can result the sample in higher hardness results [16,17]. The melt pool temperature in both delay and no delay processes were relatively constant when melt pools were formed. The melt pools of the samples deposited with no delay solidified and transformed slower as the cooling time between each laser scan was shorter, less heat conducted, the temperature of the solidified part of the sample was higher, the melt pool solidified and transformed under low temperature deference between the melt pool and the early solidified part of the sample.

In addition, the melt pools of the samples deposited with delay solidified and transformed faster as the cooling time between each laser scan was longer, more heat conducted, the temperature of the early solidified part of the sample was lower, the melt pool solidified and transformed under high temperature deference between the melt pool and the early solidified part of the sample.The melt pool firstly solidified into Austenite. Bainite can only be transformed from Austenite within certain temperature region during the solid-state phase transformation. The temperature decreasing paths of each sample were nonlinear in both delay and no delay deposition processes. The cooling time in each temperature regions could be different from sample to sample and layer to layer during each heating and cooling cycle.

The cooling time in each temperature regions could be different from sample to sample and layer to layer during each heating and cooling cycle. The different dwell time at each temperature region resulted in different percentages of Austenite were transformed463into Bainite phases during each solid-state phase transformation after the melt pool solidified into Austenite [18,19]. In this study, the factor of the samples with 10 seconds delay turned to be more durable than the samples processed with no delay can be that the time delay kept the Austenite temperature changing paths a longer time within the Bainite forming region, hence higher percentages of phase were transformed into Bainite during the solid-state phase transformation in those samples with 10 s delay.

A key stride towards the development of these systems revolves around the design of insertion-electrode batteries (IEBs). However, battery cell performance metrics of capacity and rate capability in these batteries are limited by inefficient ion and electron transport due to the complex transport channels the ions must navigate to reach storage sites – a fundamental limitation of slurry-cast (SC) type electrodes. We present a gradient-driven approach to derive optimal electrode architecture, constrained only by the underlying multiphysics system defining transport mechanisms in and across solid and liquid phases.

The derived framework challenges the traditional manufacturing techniques for electrodes, inspiring novel strategies for deriving new high-performance electrodes.Keywords: topology optimisation; nanoarchitecture; batteryIntroductionEfficient storage and deployment of renewable energy sources such as solar and wind energy are critical processes towards meeting the world's CO2 reduction targets amidst increasing energy demand. Energy demand is expected to rise by 25% by the year 2030 [1]. The key challenge revolves around the fact these sustainable sources of energy are naturally intermittent - there exist no guarantees of meeting active energy demand on-the-fly. To harness and efficiently deliver these energy sources, large-scale storage systems are being developed to ‘even-out' delivery of these energy sources. A key component of such storage systems are batteries.

A key component of such storage systems are batteries. Enhancing battery performance - capacity, rate capability, cycle efficiency and cycle life constitutes active research as scientists and engineers investigate improved battery materials and parameters [2,4].IEBs such as lithium-ion (Li-ion) and sodium-ion (Na-ion) batteries represent recent technological advances in energy storage [1]. These batteries store chemical energy and supply electrical energy via mechanisms that feature shuttling ions between electrodes and electron movement in a loaded external circuit. Detailed descriptions of IEB cell mechanisms are well- documented in literature [1, 2]. Figure 1 shows an illustration of components and mechanisms within a typical insertion-electrode cell.

Detailed descriptions of IEB cell mechanisms are well- documented in literature [1, 2]. Figure 1 shows an illustration of components and mechanisms within a typical insertion-electrode cell. Their electrochemical mechanisms depend on the ability 1815of some electrode materials to ingress or egress large quantities of lithium or sodium atoms without chemically changing the composition of the host material itself - diffusion mechanisms known as intercalation and de-intercalation respectively where atoms are inserted or removed from between the layers of the crystal structure of the host material. Traditional battery electrodes for these insertion-electrode cells consist mainly of active material particles and smaller quantities of conductive additives all held together by binding material. There exists a separation of scales of about one or two orders of magnitude between the electrode scale and the scale of the constituting particles.

There exists a separation of scales of about one or two orders of magnitude between the electrode scale and the scale of the constituting particles. As a result, battery electrodes have been historically modelled as continuum scale structures. The classical Doyle-Fuller-Newman (DFN) framework [10, 14] is the most popular physics-based continuum scale description of electrochemical processes in insertion-electrode battery cells. The framework presents a means to simulate electrochemical phenomenon efficiently and accurately without the computational intricacies of particle electrochemical kinetics.

The framework presents a means to simulate electrochemical phenomenon efficiently and accurately without the computational intricacies of particle electrochemical kinetics. Such a model accounts for the solid potential, electrolyte potential and ion concentration in the electrolyte in one spatial dimension while a radial `pseudo' dimension accounts for ion concentration in the electrode particles.Figure 1: Illustration of DFN Li-ion cell configuration showing the direction of lithium ions during a discharge, state variables, the spatial and the radial dimensions. Particle kinetics are computed from the radial dimension. Both scales are coupled by particle surface kinetics.

Particle kinetics are computed from the radial dimension. Both scales are coupled by particle surface kinetics. Binder and conductive additives have been excluded.Some works to improve cell performance have featured optimisation of cell parameters suchas separator and electrode thickness [4,5], active particle size [6] and particle distribution [7] often using a design of experiments to evaluate cell performance at discrete parameter values. This work presents a computational framework that leverages the underlying physics of ion and electron transport mechanisms to derive nanoscale topologies that optimise battery performance, specifically battery capacity and rate capability. The approach is superior to the pore engineering strategies mentioned above because it allows the physics phenomena to determine the best topology and size of architectural features of the electrode.

The approach is superior to the pore engineering strategies mentioned above because it allows the physics phenomena to determine the best topology and size of architectural features of the electrode. The framework challenges classical electrode manufacturing techniques, opening the door to other advanced manufacturing strategies for realizing the physics-1816informed electrode designs of our work.MethodContinuum scale modelConsider a DFN electrode model that features a two-dimensional spatial dimension and aradial dimension for electrode particles (such as shown in Figure 1). Spherical particles of uniform particle size have been assumed. Also, only spherical symmetrical concentration distribution is allowed and convection in the electrolyte phase is assumed negligible.

Spherical particles of uniform particle size have been assumed. Also, only spherical symmetrical concentration distribution is allowed and convection in the electrolyte phase is assumed negligible. To model effective transport properties in this continuum scale, the Bruggeman's relations have been employed so that homogenized transport properties are derived as functions of penalized bulk transport properties of solid and electrolyte phases. Diffusion within the solid particles is assumed constant regardless of degree of lithiation. No mechanical or thermal considerations have been made. For completeness, the continuum scale battery modelling equations for insertion electrode batteries are briefly enumerated below and are based on conservation of charge and mass equations in the solid electrode active material (Equations (1) and (2) respectively) and the liquid electrolyte (Equations (3) and (4) respectively).

Equation (5), known as the Butler-Volmer equation couples the system of equations across both phases by computing the lithium-ion flux across the solid-liquid interphase, ie. the solid particle surface.H∇Φs) = −asj∇. (−Ks𝜕cs 𝜕t=1 r2𝜕𝜕r(Dsr2 𝜕cs𝜕r)H [∇Φe −∇. (−Ke2RT F(1 − t+)∇ln ce]) = asjρe𝜕ce 𝜕t= ∇. (−𝐷𝑒𝐻∇ce +Ht+ 𝐢e F) +asj 𝐹j = io [exp ((1−β)Fη RT) − exp (−βFηRT)]Full derivations of these ion and electron transport equations can be found in literature [2].

Finally, the exchange current density, i0 is a function of the lithium concentration at the solid-electrolyte interface (surface of the active particles) given by:with the cell overpotential given as:βi0 = Fkcs,surf1−β(cs,max − cs,surf) ce1−βDetailed derivations of the continuum equations enumerated here are available in literature [3]. These equations are solved numerically by a time-dependent solver on a discretized domain for the solid and electrolyte phase potentials, Φs and Φe, the solid and electrolyte phase lithium-ion concentrations, cs and ce, as well as the flux transfer kinetics at the solid-electrolyte interface, j.Model EvaluationThe half-cell continuum model derived using Equations 1-7 above and parameter values ofTable 1 was evaluated by applying discharge current, ID = 30Crate A m−2 at C-rates 0.1C, 1C, 3C, 5C and 10C as shown in Figure 2. The battery cell was discharged over the time interval Ti to Tf.

The battery cell was discharged over the time interval Ti to Tf. The final time, 𝑇𝑓 was not explicitly defined but corresponded to a preset cut-off voltage of 3.33V.Battery performance was evaluated as the Energy density, W, given as:1818(6)(7)Tf W = ∫ IDVD dT Tiwhere VD represents the voltage profile during the discharge process. The results show performance peaks at different global volume fractions for each discharge rate. At low C-rates, the performance is monotonic – always increasing as the volume fraction of active material increases. However, as the discharge rate is increased, the performance is limited, reaching a peak value, and deteriorating despite further increasing active material volume fraction. Beyond the peaks, material utilization deteriorates suggesting inefficient ion transport to storage sites. To improve battery cell performance, capacity as well as rate capability must be improved.

Beyond the peaks, material utilization deteriorates suggesting inefficient ion transport to storage sites. To improve battery cell performance, capacity as well as rate capability must be improved. Active material volume fraction and indeed electrode microarchitecture is critical to cell performance. This limitation on cell performance provides inspiration for this work.This work aims to improve cell performance (capacity and rate capability), by improvingthe efficiency of ion transport mechanisms to storage sites. Rather than determining a global volume fraction, ρs which maximizes cell performance, we allow the underlying ion transport phenomenon control the volume fraction of active material locally by setting up an optimisation problem with the objective of maximizing cell performance. By determining the sensitivity of the performance metric to the local design variables, a gradient-based optimisation search is derived.

Mathematically, the problem is formulated by:1819(8)maximize: W(𝐔(ργ), ργ)subject to:F(𝐔(ργ), ργ) = 00 ≤ ∫ ργ(𝐱)dΩ ≤ ρsΩρL ≤ ργ(𝐱) ≤ ρUwhere the solution vector, 𝐔 consists of all the variables solved for as described in the continuum model formulation and ργ is the local volume fraction of active material – a function of spatial coordinates which serves as the optimisation control variables. The search is constrained by the system of equations governing ion and electron transport. The control variables are bounded locally so that the active material volume fraction lies between ρL = 0.01 and ρU = 0.9. Conductive additives and binder material is assumed to occupy 10% locally and the rest of the space is occupied by electrolyte. The global volume fraction, ρs = 0.6 is set so that at optimisation initialization, ργ is 0.6 everywhere. The global volume fraction is constrained during the optimisation.

The global volume fraction, ρs = 0.6 is set so that at optimisation initialization, ργ is 0.6 everywhere. The global volume fraction is constrained during the optimisation. The real-valued objective functional, W, depends on the solution vector, 𝑼. It also depends on the vector of control variables, 𝜌𝛾. Sensitivities of the objective to the control variables have been determined by the adjoint method. Optimisation termination occurs when the objective value converges within a tolerance of 10−4 or the maximum number of iterations is reached.Results and DiscussionFigure 3 shows the optimal layout of the electrode active material volume fraction at optimisation termination. The solution shows that cell performance is maximized when ion and electron transport is facilitated to all regions of the domain.

The solution shows that cell performance is maximized when ion and electron transport is facilitated to all regions of the domain. Channels consisting of electrolyte project into the electrode domain towards to the current collector side, creating layers of dense electrode material.ρ𝛾Figure 3: Optimised electrode topology described by optimal layout of the local volume fraction of active materials. Dark regions are predominantly electrolyte and the white regions indicate active material of highest volume fraction. The grey regions are regions of porous electrode material with intermediate volume fraction.The paths taken by Li-ions into the electrode are plotted in the streamlines of Figure 4. Regions of tightly packed streamlines indicate preference for electrolyte phase transport as the Li-ions travel to regions of previously poorly utilized active material.

Regions of tightly packed streamlines indicate preference for electrolyte phase transport as the Li-ions travel to regions of previously poorly utilized active material. With the homogeneous SC type electrode, these regions were difficult to be reached by ions due to the effective tortuous random channels available for electrolyte percolation. The streams lines can be observed to bend as the Li-ions intercalate into the active materials layers.1820Figure 4: Lithium-ion flux streamlines, originating from the negative electrode surface, travel across the separator into the optimised continuum layout of active materialsThe optimised electrode also offers an increased depth of discharge at cutoff voltage which terminates the cell discharge as shown in Figure 6a.

At 5C discharge rate, the optimised electrode offers 15.4% increase in areal capacity compared to the uniform electrode at cutoff voltage.(a)(b)Figure 5: (a) Depth of discharge into positive electrode at 5C discharge rate for SC and Optimised electrodes. (b) Energy density at different C-rates for SC and optimised electrodes, showing performance increase at different C-rates (green dotted markers).The improvement in cell performance is shown in Figure 6(b) where the variation of energy density achieved at discharge termination is compared for both electrodes for a range of discharge rates. It can be observed that there is a 22.4% increase in energy density when the optimised electrode is discharged over the uniform electrode at 5C discharge rate. The variation of gains in performance for different C-rates also shows that at low C-rates (below 1C), the optimised electrode offered no significant advantage in performance.

The variation of gains in performance for different C-rates also shows that at low C-rates (below 1C), the optimised electrode offered no significant advantage in performance. This is consistent with observations of other works since discharge at low current density provides sufficient time for lithiation throughout the electrode domain, regardless of electrode architecture. Hence, the benefits of an optimised electrode topology increase with discharge rates as uniform electrodes suffer deteriorating rate capability and material utilization. In the range 3C to 5C, the percentage increase in performance is most sensitive to C-rate suggesting a practical range for use during gradient-based optimisation formulations involving time- dependent discharge process.1821Summary and ConclusionA continuum scale P3D model was developed according to established continuum equationsdescribing ion and electron transport in and across solid and electrolyte phases.

The model was evaluated at a range of C-rates and active material volume fractions to highlight the significance of active material layout to cell performance. By mathematically formulating an optimisation problem, the energy density of the electrode was maximized by controlling local volume fraction of the electrode material. The result was the generation of optimal layout of electrode material for maximizing the cell performance. The optimal topology was shown to facilitate ion transport to hitherto under-utilized regions of the electrode domain. Most importantly, the optimised electrode offered over 15% increase in areal capacity and over 22% increase in areal energy density when discharged at 5C discharge rate. Further improvements in performance were attained for higher C- rates.

Further improvements in performance were attained for higher C- rates. This work demonstrates that physics-informed computational optimisation can identify optimal electrode topology for efficient ion and electron transport, maximizing battery performance.In future works, model fidelity will be enhanced to support a more robust cell performanceoptimisation by coupling diffusion and stress to state of charge. Proposed architectures will be tested experimentally using these patterned electrodes.Acknowledgement The authors acknowledge the funding support of Shell and the contributions of Peter Klusener, Magda Titirici, Ifan Stephens and all other collaborators on the NanoBat project.References[1]Plett G.L. Battery Management Systems: Battery Modelling, 2015[2]Gupta A., Seo J.H., Zhang X., Du W., Sastry M.A., Shyy W. Effective transport properties of Lithium Manganese Oxide Electrode via Particle-Scale Modeling.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDiscrete Element Modeling of Fused Deposition Modeling ProcessMenezes, C. and Turner, C.*Department of Mechanical Engineering, Clemson University, Clemson, SC, USA 29631 * Contact Author (cturne9@clemson.edu)AbstractFused Deposition Modeling components show anisotropic material properties as a result of the difference in bonding strengths between inter- and intra-layer particles. This difference occurs due to temperature gradient of the deposited filaments which affects the filament adhesion as well as due to the presence of voids or other discontinuities in the printed object. Discrete Element Modelling (DEM) is a discontinuous methodology which follows the idea of treating filaments as discrete rigid particles with simplified geometries for calculating the thermal interactions between the particles.

Models using this approach allow the investigators to correlate the adhesion effects between filaments based on experimental data which ultimately will allow for optimization of the relevant printing parameters.IntroductionAdditive Manufacturing (AM) builds parts by sequentially adding material, often in the form of layers to an existing part, as opposed to the traditional practice of subtractive manufacturing where material is removed from a larger piece of material to leave a desired part. Multiple types of AM processes exist, including Stereolithography (SLA), Selective Laser Sintering (SLS), Fused Deposition Modelling (FDM), Inkjet Modeling (IJM), Direct Metal Deposition (DMD), and Layered Object Manufacturing (LOM).

The differences in these techniques are largely defined by the mechanism by which new layers are created and by the materials usable in each technique [1].While the initial machines were expensive and thus limited the technology to specialty customers [2], technology development has enabled these machines to move into increased use for prototyping and now for manufacturing. Consumer grade versions of many of these machines are already available and support a growing competitive home-market for AM products as well as a competitive sector of companies focused on “job-shop” AM services [3]. FDM is one of the easiest technologies to commercialize and is often an entry point for AM systems. FDM is primarily a polymer-based AM technique [2] that extrudes raw material in filaments through a heated extrusion nozzle as shown in Figure 1.

FDM is primarily a polymer-based AM technique [2] that extrudes raw material in filaments through a heated extrusion nozzle as shown in Figure 1. Filaments are extruded in adjacent passes of the machine to form a layer, and subsequent layers are built one on top of another.The quality and material properties of FDM parts are effected by various in-process parameters such as layer thickness, build orientation, print speed, extrusion temperature, infill density, and infill pattern [4], in addition to other parameters. Additives used to color or otherwise modify the filaments can decrease the quality and mechanical performance of FDM manufactured parts [5, 6] and environmental factors like humidity and temperature [6] can also affect the quality of the builds. Part quality is also dependent upon the accuracy and precision of the slicer software1334that converts the CAD model to .STL file and the .STL file to g-code instructions which is the tool path followed by the nozzle.

Machine parameters may further limit the quality of the build by the introduction of dynamic effects. The print time leads to different bonding behaviors when comparing the bonds within a single filament (intrafilament bonds), between filaments within a layer (intralayer bonds), and between layers (interlayer bonds). Cumulatively, these effects lead to FDM manufactured parts that may exhibit non-homogeneous, anisotropic and nonlinear behaviors [7], and to the products have quality control issues including dimensional errors, delamination between filaments and layers, undesirable porosity and poor or unpredicted material properties [8].

Yet, the ability to build simulations of the build process [3] that can lead to models that can accurately and reliably predict the material properties of the finished product in needed [9].FIGURE 1: Schematic of the FDM Process, Adapted from [4].BackgroundFollowing a path determined by the slicing software, FDM moves the extruder through the workspace depositing semi-molten material to form filaments. The bonds formed by the filament as it cools are primarily driven by the initial thermal energy of the semi-molten material at deposition [10] and the resulting polymer diffusion enabled by the filament temperature [11]. Cooling of the filament is controlled by heat transfer to the ambient environment (primarily through conduction and convection to the air, by conduction to adjacent materials in contact with the filament (i.e. other filaments and the build plate) [12].

other filaments and the build plate) [12]. The cooling profile of the filament as a function of its position and time is therefore an important factor in the resulting mechanical properties of the printed part.Prior research has considered this problem through the use of finite difference methods [13, 14] and finite elemental analysis [15-17]. These methods assume that the part is a continuum. Other researchers, [12, 18, 19] have been developed analytical solutions where the transient heat model includes activating or deactivating all relevant local boundary conditions depending on part1335geometry, operating conditions and deposition strategy. A related mathematical approach treats the filament as discrete particles (or elements) and can be used to construct systems with various boundary conditions to simulate thermal and mechanical changes between elements [20].

This discrete element method (DEM) approach is particularly popular when simulating powder-based AM techniques such as SLS [21]. Our focus is in developing these methods as a means to simulate the behavior of an FDM process by representing the filament as a sequence of discrete particle elements.MethodologyIn a typical DEM approach, a discrete element of spherical shape is used to represent an element. A spherical shape is used because it simplifies the contact calculations with adjacent elements. However, when simulating a filament, the initial spherical shape quickly “slumps” to form a filament with an oblong cross-section. While a higher fidelity element model could be used to account for this mechanical slumping behavior, we chose in our initial model to simply assume that the element remains round and will have a diameter equal to the layer height of the printed part. Consequently, the filament width is equal to the layer height.

Consequently, the filament width is equal to the layer height. Note that this assumption may prove to be inaccurate, but for an initial model, we feel that it is a reasonable first approximation of the behavior of the filament.Furthermore, for our initial model, we decoupled the g-code process from the model creation process and focused on the development of rectangular parts with a solid infill pattern composed of filaments deposited as shown in Figure 2. Since for most FDM printers, the layer height values range from (0.2 mm to 0.4mm), the resulting Biot number (Bi), is less than 0.1. Hence, each element can be considered to be a lumped capacity element which assumes uniform temperature distribution across the cross-sectional element area. Thus, each layer is composed of a single filament (18 elements long in Figure 2) laid down in a zig-zag pattern. As is typical of slicing programs, subsequent layers are often formed with a zig-zag patten orthogonal to the layer below.

As is typical of slicing programs, subsequent layers are often formed with a zig-zag patten orthogonal to the layer below. We adopted this approach to simplify the initial program and will integrate a g-code based path planning algorithm in future versions of the code.a. First Odd Layerb. First Even LayerFigure 2: Modeled Deposition Sequence.1336Based on these assumptions, we can define some of the mathematical properties of theelements.

First Odd Layerb. First Even LayerFigure 2: Modeled Deposition Sequence.1336Based on these assumptions, we can define some of the mathematical properties of theelements. The surface area of the element is that of a sphere with a radius, R.24AR=Similarly, the element has a volume of a sphere of radius, R.343VR=Since the thermal energy stored in an element is a function of the volume, density, heatcapacity and the temperature of the element we can write()depositionpdepositionambUVCTT=−Here, we make another assumption, that the resulting element has “good” contact with anyadjacent elements such that the area in contact is given by6contactAA=While a true rigid sphere would have only point contact with its neighbor, this contact approximation is that of a cube in contact with other cubes. Since in reality our sphere is not rigid, but is in fact “mushy”, this assumption is probably closer to the reality, although it is not also imperfect.

Since in reality our sphere is not rigid, but is in fact “mushy”, this assumption is probably closer to the reality, although it is not also imperfect. Thus, by applying Fourier’s Law, we can find that the conductive heat transfer is()ABCthTTQR−=where the thermal resistance Rth isthcontactLRkA=where L is the characteristic length, and k is the material conductivity. For this model, the characteristic length is given by3contactVRLA==We can apply similar techniques to model convection and radiation effects on each of the six faces of the element. Initially, we will assume that only free convection is present, and that radiative heat transfer is negligible.

Initially, we will assume that only free convection is present, and that radiative heat transfer is negligible. Thus, we can determine that the internal energy, Ui, of the element at a time ti, is given by:1337(1)(2)(3)(4)(5)(6)(7)11iiconductionconvectioniiUUQtQtandttt−−=−−=−and so, the temperature Ti, of the element is given by rearranging Equation (3) asiiambpUTTCV=+where we assume that the deposition time, ti, of the element is at i = 0. At deposition, the temperature of the element is that of the extruder, which we refer to as the deposition temperature.Based on this formulation, we can have several different heat transfer boundary conditions that need to be modeled. The first element deposited (element 1 in Figure 2) has five faces exposed to the ambient environment, where conduction to the ambient atmosphere and free convection are the modes of heat transfer. These conditions apply to the left, right, front, back and top faces of the element.

These conditions apply to the left, right, front, back and top faces of the element. However, the bottom face is in “good” contact with the build plate and in this case only experiences a conductive heat transfer to the build plate. If the build plate is cooler than the element, heat flows from the element into the build plate, cooling the element, and if the build plate is warmer than the element, heat flows from the build plate into the element, further warming the element.When the second element is deposited next to element 1 (on the right side of element 1) in Figure 2, the right hand side of element 1 changes its boundary conditions. Element 1 no longer has a free convection mode on its right face, and its conduction mode changes from the ambient atmosphere to a conduction mode with element 2. The other faces remain in the same heat transfer states.

The other faces remain in the same heat transfer states. Element 2 begins with free convection and atmospheric conduction on four faces, right, front, back and top, but experiences conduction with element 1 on its left face, and conduction with the build plate on its bottom face.Ultimately, there are 12 heat transfer cases that need to be considered. Six of the cases involve the build plate and various numbers of faces exposed to the atmosphere or to other elements. The remaining 6 cases involve situations where some faces are in contact with other elements and some faces are in contact with the ambient atmosphere. The model determines the appropriate boundary case for each element based on the flowchart presented in Figure 3.The model was initially formulated in MatLABTM to evaluate our results against several test cases to determine if the initial assumptions appear to be valid.

This formulation would also allow us to use the model to compare to previous experimental data sets generated by [9]. However, due to the larger size of these data sets, use of the campus Palmetto cluster would be necessary. Fortunately, with the model formulated in MatLABTM, the process of running the model on the cluster did not require substantial modifications to the existing code.1338(8)(9)FIGURE 3: Model Flow Chart.Preliminary Validation TestsFor this work, we will present the results for three test cases representing specimens of different sizes. The process parameters and materials properties for the tests are shown in Table 1. The parameters of the three test cases are shown in Table 2.1339Table 1.

Odd layers are thus printed in a horizontal direction while even layers are printed in a vertical direction. The third data set is also rectangular but printed in a 90 degree orientation to that of dataset 1, and comprises ten layers instead of six. Again, odd layers are printed horizontally, and even layers are printed vertically. For each test case, we recorded the temperature histories for each element. Figure 4 shows the temperature history for all three datasets at the midpoint of the printing process. For dataset 1 that would be after the 750th element, for dataset 2 the 768th element, and for dataset 3 after the 1250th element. Elements not yet printed show a printing temperature of 0 Kelvin in Figure 4.FIGURE 4: Element temperatures at printing midpoint.1340Note in Figure 4 that the most recent element is also the element at the highest temperature and that the elements are generally cooler the longer it has been since the element was first deposited.

It is also apparent that the cooling rate is not uniform, and in fact we see some variation of the temperature of the elements with the elements on the edges of the specimen cooling faster than those internal to the specimen. This phenomenon is more extreme when the printing direction is a shorter length as is the case with the odd layers in dataset 3 and the even layers of dataset 1 (where the filament length is each 10 elements) compared to the longer 16 element lengths in dataset 2 and the 25 element lengths in datasets 1 and 3. This makes physical sense in that the external elements are better able to expel heat than the internal elements.If we examine the same data once the entire part has been printed, as shown in Figure 5, we see the same trends are also apparent.

Notice that the elements in the part now have temperatures ranging from about 340K to the extruder temperature of 493K.FIGURE 5: Element temperatures at printing endpoint.Once the print is complete, we begin a process of allowing the build plate to cool from a heated temperature of 333K to a plate temperature 298K over several minutes. We continue to monitor the temperatures of the elements awaiting the part to cool to ambient temperature. Figure 6 shows the temperature profiles a quarter of the time through this cooling process. Note that the cooling pattern has inverted itself, and the elements farthest from the build plate are now the coolest elements in the model. Similarly, the span of temperature of the part is now less than 3K as opposed to the 160K when the build was complete.

Similarly, the span of temperature of the part is now less than 3K as opposed to the 160K when the build was complete. And while the edge elements are slightly cooler than the interior, the temperature gradient between the interior and edges is far less significant than it was earlier in the simulation.1341FIGURE 6: Element temperatures at a quarter of the time through the cooling process.In Figure 7, the temperature profiles of the elements at the midpoint of the plate coolingprocess continue to show a similar trend.

Except now the range of temperatures is about 1.5K.FIGURE 7: Element temperatures at half of the time through the cooling process.While Figure 8 initially appears to show that the temperature gradient within he part is increasing, by examining the temperature axis, it can bee seen that the range of variation is now approximately 0.5K, indicating that the part is continuing to approach an isothermal state.1342FIGURE 8: Element temperatures at three-fourths of the time through the cooling process.Finally, in Figure 9, we observe that the part has effectively cooled to the build platetemperature of 298K, with a temperature variation of approximately 0.5K throughout the part.FIGURE 9: Element temperatures at the end of the simulation.These behaviors appear to mimic observed physical processes, where the part cools from the outside-to-the-inside, and where heating the build plate prolongs the time that the layers closest to the build plate build polymer connections between filaments (intralayer) and between layers (interlayer).

This further explains the anisotropic behaviors observed, where the layers adjacent to the build plate approximate bulk material properties, while the layers furthest removed from the build plate demonstrate increased anisotropic behaviors related to the filament orientations.1343Conclusions and Future WorkThese trials demonstrate that the DEM simulation constructed appears to mimic the thermal behavior of cooling FDM materials despite the simplifying assumptions made to construct the DEM model. Encouragingly, we even observed cooling differences due to the proximity to the edges of the build and due to the warming effects of a heated build plate. We believe that qualitatively, the model is matching observed behaviors.However, quantitative evaluation of the model remains incomplete. We will need to evaluate the observed cooling rates versus experimental data to confirm that the predicted cooling temperatures are valid.

We will need to evaluate the observed cooling rates versus experimental data to confirm that the predicted cooling temperatures are valid. In addition, the incorporation of a g-code based scheme to match a generically sliced model is necessary to compare to the much wider range of print configurations available in actual AM processes. Furthermore, there are several opportunities to improve the computational performance of the model.Our initial goal is to analyze the thermal behavior of elements in the model to predict the time during which adjacent elements are at a suitable temperature to form bonds between the elements. This relationship should demonstrate that there are bond formation time differences between intrafilament, intralayer and interlayer bonds that correspond to observed differences in material modulus in these directions [9].

This analysis should assist in determining whether a simple thermal model is adequate, or whether a more complex mode incorporating surface effects is necessary to predict these bond strengths within a desired range of accuracy.Additional effects, such as the mechanical “slump” experienced by the filament upon deposition (attributable to the viscosity of the material) and as a result of the weight of additional layers added on top of the initial layers of filament should allow for the model to also account for dimensional effects furing printing, and potentially even curling effects. Understanding these effects in simulation may allow for advanced build plates that allow for regionalized heating and cooling of the part during the build process to optimize dimensional stability and promote bonding.

Furthermore, future DEM models can use the build information to account for induced stresses and to predict part strengths and anisotropic properties by incorporating additional models of multiphysics effects. This capability would allow a part to be optimized with the design-build-and- use cases fully considered before the part is ever printed, supporting process certification.The initial model described in this work is but a first step towards a full multiphysics model that can examine a host of potential improvements to the FDM process. Furthermore, the insights gained in developing this model can be applied to models of other AM processes leading to a fuller picture of the complex multiphysics effects present in AM processes. Understanding, predicting, and using these effects is entirely within the realm of possibilities for computationally supported AM models as a result of this work and work like this.

Understanding, predicting, and using these effects is entirely within the realm of possibilities for computationally supported AM models as a result of this work and work like this. We are pleased with our initial results and encourage work to continue in this and other avenues to support a multiphysics understanding of additive manufacturing from design, through the build process, to the end use of the parts.1344AcknowledgementsThe authors would like to acknowledge the support of the Department of Mechanical Engineering at Clemson University for their support of this work. All statements within are those of the authors and may or may not represent the views of these institutions.References[1][2] [3][4][5][6][7][8][9][10][11][12][13][14][15][16][17][18][19]Pereira, T., Kennedy, J. V., and Potgieter, J., 2019, "A Comparison of Traditional Manufacturing Vs Additive Manufacturing, the Best Method for the Job," Procedia Manufacturing, 30 pp.

V., and Potgieter, J., 2019, "A Comparison of Traditional Manufacturing Vs Additive Manufacturing, the Best Method for the Job," Procedia Manufacturing, 30 pp. 11-18. Rafiq, N., 2006, " Rapid prototyping : principles and applications," Wiley, Hoboken, N.J. Migler, K., and Ricker, R., 2016, "Measurement Science Roadmap for Polymer-Based Additive Manufacturing," Advanced Manufacturing Series, NIST ADM, https://doi.org/10.6028/NIST.AMS.100-5 . Dey, A., and Yodo, N., 2019, "A Systematic Survey of FDM Process Parameter Optimization and their Influence on Part Characteristics," Journal of Manufacturing and Materials Processing, 3 pp. 64. Pandžić, A., Hodzic, D., and Milovanović, A., 2019, "Influence of Material Colour on Mechanical Properties of PLA Material in FDM Technology ," . Valerga, A. P., Batista, M., Salguero, J., 2018, "Influence of PLA Filament Conditions on Characteristics of FDM Parts," Materials (Basel, Switzerland), 11(8) pp. 1322.

T., 2003, "Experimental Study of the Cooling Characteristics of Polymer Filaments in FDM and Impact on the Mesostructures and Properties of Prototypes," Solid Free.Fabr.Proc., Austin, Texas, pp. 313-323. Coogan, T. J., and Kazmer, D. O., 2017, "Healing Simulation for Bond Strength Prediction of FDM," Rapid Prototyping Journal, 23(3) pp. 551-561. Costa, Duarte, F., and Covas, J., 2015, "Thermal Conditions Affecting Heat Transfer in FDM/FFE: A Contribution Towards the Numerical Modelling of the Process," Virtual and Physical Prototyping, 10. Zhang, J., Wang, X. Z., Yu, W. W., 2017, "Numerical Investigation of the Influence of Process Conditions on the Temperature Variation in Fused Deposition Modeling," Materials and Design, pp. 59-68. Stockman, T., Schneider, J., Walker, B., 2019, "A 3D Finite Difference Thermal Model Tailored for Additive Manufacturing," Jom, 71.

59-68. Stockman, T., Schneider, J., Walker, B., 2019, "A 3D Finite Difference Thermal Model Tailored for Additive Manufacturing," Jom, 71. Ji, L., and Zhou, T., 2010, "Finite Element Simulation of Temperature Field in Fused Deposition Modeling," Advanced Materials Research, 97-101 pp. 2585. D’Amico, A., and Peterson, A. M., 2018, "An Adaptable FEA Simulation of Material Extrusion Additive Manufacturing Heat Transfer in 3D," Additive Manufacturing, 21 pp. 422-430. Zhou, Y., Lu, H., Wang, G., 2020, "Voxelization Modelling Based Finite Element Simulation and Process Parameter Optimization for Fused Filament Fabrication," Materials & Design, 187 pp. 108409. Costa, S., Duarte, F., and Covas, J., 2017, "Estimation of Filament Temperature and Adhesion Development in Fused Deposition Techniques," Journal of Materials Processing Technology, 245 pp. 167- 179.

screte Multi-Material Selective Laser Sintering (M2SLS): Development for an Application in Complex Sand Casting Core ArraysBrad Jackson, Kris Wood, Joseph J. BeamanDepartment of Mechanical Engineering The University of Texas at Austin Austin, TX 78712AbstractConventional sand casting processes often take many weeks or months to produce thetooling required for a casting, in particular the fabrication of sand core arrays for hollow features in a casting. SLS is already being applied to produce complex sand core geometries and reduce production times, but a new development of discretely laying down two different materials and removing one after sintering will allow even more complex geometries and drastically decrease the production times of sand cores. Two of the most significant problems in the current use of SLS for sand cores are the mechanical removal of unsintered powder and damage during part breakout.

Two of the most significant problems in the current use of SLS for sand cores are the mechanical removal of unsintered powder and damage during part breakout. The second discrete material serves as a support medium through the build and fabrication of the sand core and is removed before casting; the sacrificial second material increases green strength and eliminates time consuming post-processing. The development and plan for implementation of the discrete M2SLS process is presented.IntroductionSince the introduction of Selective Laser Sintering (SLS) at the University of Texas atAustin and other solid freeform fabrication, the evolution from a single material to multiple materials was only a matter of time. Currently, many researchers are pursuing multiple material solid freeform fabrication processes [1,2]. Multiple material SLS (M2SLS) can be divided into two categories: discrete and functionally gradient.

Multiple material SLS (M2SLS) can be divided into two categories: discrete and functionally gradient. Discrete refers to parts built with distinct material regions and interfaces within the part; each material region consists of one material composition. Functionally graded are parts that the material composition within the part changes from one component to another over a finite distance. The grading or blending is functional because the grade allows for the tailoring of physical, electrical, or chemical properties within a single part. Current research at the University of Texas at Austin examine functional graded parts with a WC-Co material system where the percent weight composition of WC is varied to achieve desirable mechanical properties at specific regions [3].

In this paper, the development pertains to a discrete two material system.Sand casting is an old technology, dating back a thousands years, but still is the mostprevalent form of casting, with 15 million tons of metal cast in the United States each year [4]. Simply put, sand casting is the process of imprinting a pattern into sand to form a mold, placing cores into the mold for internal cavities, including a gating system for molten metal, pouring the molten metal, and after cooling, breaking the mold to remove the casting. Sand castings allow for complicated shapes at a relatively low cost. The introduction of rapid prototyping to sand casting is fairly new but has impacted the industry with reduced production times and increased176complexity not possible without SFF [5]. There are two RP implementations in sand casting: pattern making and coremaking.

There are two RP implementations in sand casting: pattern making and coremaking. Patterns are used to form the imprint in the sand mold.Figure 1: Core array and finished castingCore arrays form the internal cavities or passageways of casted metal parts, such as internal geometries of engine blocks. Figure 1 shows a complex core array and the resulting finish casting. In conventional coremaking, cores are fabricated in many pieces that would have to be assembled with core paste and placed in the mold and gauged for accuracy of the fit. Cored holes and features are usually limited to ¼" diameter, but there is one special technique that exists to produce small passageways in sand castings; it is called the Dalton® Core. The process uses a copper tubing that is bent to the desired shape and inserted in the mold. A braided refractory sleeve of glass or stainless steel covers the outside of the tube.

The process uses a copper tubing that is bent to the desired shape and inserted in the mold. A braided refractory sleeve of glass or stainless steel covers the outside of the tube. A chemical solvent, which aluminum is insoluble, removes the copper tube, while the sleeve remains. The sleeve is then removes manually from the hole [6]. This special technique has very limited application, for it cannot be used for varying hole size or complex contoured shapes.The SLS rapid prototyping process is being used to produce sand casting core arrays.Sandform™, introduced by DTM Corporation, is a polymer-coated sand that can be processed to produce single complex core arrays, as shown in Figure 1. With the implementation of Sandform, the whole core array can be manufactured in a SLS part, with no core assembling and reduced gauging. Figure 2 shows the coremaking procedure using Sandform and SLS.

Figure 2 shows the coremaking procedure using Sandform and SLS. This new technology was hailed by the casting industry, but the process has many problems and limitations.177Shell Core - 1/8" – 3/8" thicknessPre ProcessesAdd furniture - support features - bone structureCreate .stl files -fill scan -outline scan (high power)Pre CureBreakout in process chamberClean with soft brushes and vacuumPost ProcessesRepair fractures and damagesFlame cure surfacePlace in curing bin with mediacure: 170 C to 180 C for 4 hours (Solidiform) 160 C for 4 hours (DTM)Repair post cure damageEliminated by using second materialFigure 2: Sandform Process178Temperature -ramp 70C to 60C over .5 in. -feed heater 50 CCurrent problems with SandformBecause cores are produced by SLS, extra polymer binder must be added to the sand/binder system to achieve a satisfactory green strength for the SLS build and handling of the part.

This added binder causes more outgassing during metal pouring and solidification. With added binder and therefore more outgassing, venting of the core arrays become more important to prevent gases entering the molten metal and forming pockets or cracks.Venting is achieved by creating passageways inside the core itself. Figure 3 shows aCAD representation of a core with the necessary venting. The vents connect to a core print in the mold to discharge the gas to the atmosphere. For any sand core feature, a vent hole must be present to outgas steam and gas from the polymer degradation during metal pouring; thus for a .125" cylindrical feature, an internal hole must be present in the feature to allow proper venting and still have strength.

The vents inside the core can become quite small and intricate, and as mentioned before are limited to ¼" diameter, which restricts the feature sizes that can be realized.Another problem related to venting is the mechanical removal of the support material orthe unsintered powder. The problem actually encompasses all SLS manufactured parts with small holes or cavities. The unsintered powder must be mechanically removed from the cavities. Because of the vents in the sand cores, the sand cores are basically shelled, so they are filled with unsintered sand powder. The low green strength of Sandform adds to the problem of removal because of the increased likelihood of breakage. Common removal methods are gravity, shaking or vibration, and pressure differences (suction or blowing), but with smaller and more complex features, the above methods cannot remove the powder [5].

Common removal methods are gravity, shaking or vibration, and pressure differences (suction or blowing), but with smaller and more complex features, the above methods cannot remove the powder [5]. Figure 3 is an example of a core where the unsintered powder could not be removed.Figure 3: CAD core and section view179In addition, low green strength can cause part damage during breakout of the SLS machine and removing the unsintered powder and part cleanup. Some damage can be repaired with special adhesives, but many cores have to be scrapped. Also, the core can break or crack during the thermal cure cycle. As stated before, binder content cannot be increased to augment the green strength because more binder means more outgassing and more need for venting or larger vents.All of these problems point to the overlying problem of the post-processing time of theSandform sand core arrays.

Between cleaning and repairing, a person can spend over four hours per part, which is a considerable amount of time even in low volume runs of less than a 100 pieces [5]. Granted, the SLS sand cores do cut the production times of cores, but there is still an incredible potential to improve the process.A two material discrete system solves the problems and improves processSeeing the current problems with Sandform and realizing the possibilities, one would think this a perfect opportunity to employ a multiple material system. A two material SLS system would greatly improve the functionality of Sandform. First, the unsintered powder removal step would be eliminated from the process, cutting the post-processing time; at the same time, handling would be minimized because there is no need to handle the core to remove loose powder, which would also cut down on the probability of breakage and therefore repairing a damaged core.

Next, the second material would add green strength to the core, acting as a skeleton to give the part strength and support, which means less breakage.Besides solving the current problems, a two material system would enhance Sandform.The hassle of removing unsintered powder would be eliminated to allow smaller and more complicated features to be implemented in core designs. Small features require small venting, so previous designs that were impossible to make because of the difficulty of removing unsintered powder can be made possible. Also, the added strength and support of the second material would give the necessary integrity for smaller and more complex features. Another added benefit is in the mold assembly.

Also, the added strength and support of the second material would give the necessary integrity for smaller and more complex features. Another added benefit is in the mold assembly. Since the core can be made in a single piece, there would be less time needed to fit the core to the core print, and better accuracy could be achieved because there would be less chance for error in the placement of the core.Concept developmentIn order to develop a discrete multiple material system in a timely and effective manner, some limitations were imposed on the project from the outset. The system is to work within the Sandform coremaking process parameters. The research is driven by the application to sand cores; therefore, the overall goal is to produce complex sand core arrays with Sandform, not necessarily to produce a whole new system.

The research is driven by the application to sand cores; therefore, the overall goal is to produce complex sand core arrays with Sandform, not necessarily to produce a whole new system. This is not to say that the discrete M2SLS system could not be applied to other applications or is not designed with other applications in mind, but the focus is Sandform.Having defined the goal, the development work was decomposed into three main functions: powder delivery, material selection, and powder removal. For each function,180brainstorming and other concept generation techniques were engaged to find possible solutions. Then, the possible solutions or ideas were weighed against a criteria, including feasibility, complexity, and cost. From the criteria, concepts were combined for each function.

Then, the possible solutions or ideas were weighed against a criteria, including feasibility, complexity, and cost. From the criteria, concepts were combined for each function. Although material choice and powder removal are independent functions, the two are actually coupled to each other in the synthesis of a concept because the removal of powder is process dependent which would affect the material selection.Because of intellectual property issues at this time, details cannot be discussed in this paper concerning the concept development or prototyping of the discrete multiple material system; however, the criteria for development is given below:Material Selection and RemovalDoes not interfere with sand/phenolic system • Minimize residue • SLS “friendly” material • Fast removal • Inexpensive • Safety and minimize health risks Powder DeliveryMinimize mixing between layers • Use SLS technology • SLS “friendly” powders • Minimize complexity and control of system (K.I.S.S.

principle) • Adaptable to current SLS machinesMany materials were tested for use as the secondary discrete material. Tests included thermogravimetric analysis (TGA), as well as performance in the SLS process. The tests yielded two to three possible materials for further experiments with the entire prototyped system. A proof-of-concept prototype was built and tested the feasibility of the powder delivery system. It exhibited favorable results with many insightful solutions to problems that were unforeseen.Future work and conclusionsThe next step is to create a beta prototype that incorporates the enhanced powder delivery system and the use of possible materials. This system will produce simple shapes and parts that can be further tested in the Sandform coremaking process, along with more extensive material testing with the Sandform and second material.

This system will produce simple shapes and parts that can be further tested in the Sandform coremaking process, along with more extensive material testing with the Sandform and second material. Once the material system and powder delivery system have been proven satisfactory, foundry tests will be performed with simple shapes to verify the workability of the system with foundry standards.This paper has presented the advantages, problems, and potential of Sandform in the sand casting industry. Sandform is successful but has its problems and limitations that could be solved and181enhanced with the development of a discrete multiple material SLS system. Since a solid design approach was followed, the prototypes have produced promising results, and the material system has encouraging possibilities.References[1] Rock S.J. and C.R.

Since a solid design approach was followed, the prototypes have produced promising results, and the material system has encouraging possibilities.References[1] Rock S.J. and C.R. Gilman, "A New SFF Process for Functional Part Rapid Prototypingand Manufacturing: Freeform Powder Molding," Proceedings of the Solid Freeform Fabrication Symposium, The University of Texas at Austin, August, 1995.[2] Kumar, Ashok and Hongxin Zhang, "Electrophotographic Powder Deposition forFreeform Fabrication," Proceedings of the Solid Freeform Fabrication Symposium, The University of Texas at Austin, August 1999.[3]Jepson, L. et al, "Multi-Material Selective Laser Sintering: Empirical Studies and Hardware Development," Proceedings of the 2000 NSF Design and Manufacturing Grantees Conference, January 2000.[4] Kalpakjain, Serope, Manufacturing Engineering and Technology (Third Edition),Addison-Wesley, NY, 1995.[5] Andre, Larry Sr. and Harold Luper, personal interviews, Solidiform, Inc.

SLOCATION DENSITY BASED FINITE ELEMENT MODELING OF ULTRASONIC CONSOLIDATION D. Pal and B. E. Stucker, Utah State University, Logan, UT 84341AbstractA dislocation density based constitutive model has been developed and implemented into a crystal plasticity quasi‐static finite element framework. This approach captures the statistical evolution of dislocation structures and grain fragmentation at the bonding interface when sufficient boundary conditions pertaining to the Ultrasonic Consolidation process are prescribed. Hardening is incorporated using statistically stored and geometrically necessary dislocation densities (SSDs and GNDs), which are dislocation analogs of isotropic and kinematic hardening respectively. The GND considers strain‐gradient and thus renders the model size‐dependent.

The GND considers strain‐gradient and thus renders the model size‐dependent. The model is calibrated using experimental data from published refereed literature and then validated for the Aluminum 3003 alloy.IntroductionAs a direct result of ongoing research efforts in ultrasonic consolidation (UC) worldwide [1], it has become apparent that a new approach to modeling of UC bonding is needed. A model which provides a better understanding of the effects of process parameter changes on grain refinement, plastic deformation and bonding during UC will better enable researchers to predict which materials will bond, how the mechanical properties of UC-produced parts can be improved, and how to better design the next generation of UC equipment.The continuum properties of parts made using UC are strongly dependent upon the micromechanics of the bonded interface [1].

Interfacial-scale microstructures can be studied fundamentally using electron microscopy [1] and can be used to correlate atomic and mesoscopic mechanisms of deformation to their continuum counterparts. A dislocation density-based crystal plasticity Finite Element Model (FEM) can capture the statistical distribution of dislocations, partials and various deformation mechanisms at the bonding interface as inputs to predict macroscopic deformation profiles as a function of energy input characteristics. These input characteristics are a function of the process parameters used in a UC machine, namely vibration amplitude, normal force, ultrasonic frequency, welding speed, sonotrode geometry and temperature.Problem FormulationIt has been shown that material sheets subjected to UC undergo inhomogeneous plastic deformation through their thickness [1]. Classical continuum plasticity theories do not fully explain this phenomenon [2].

Classical continuum plasticity theories do not fully explain this phenomenon [2]. Therefore, a study of strain localization and grain refinement at the material interfaces during UC bonding is required. The following steps lead to the calculation of these localized strains and their effects.Large Deformation Quasi-Static Crystal Plasticity DescriptionThe deformation map in space and time is described by the total deformation gradient tensor F (Figure 1). Applying the Kroner-Lee assumption, F is decomposed into elastic Fe and plastic gradient Fp tensors using multiplicative operator theory(Eqn 1)(cid:1) (cid:2) (cid:1)(cid:3)(cid:1)(cid:4)Figure1 Multiplicative decomposition of the total deformation gradient, F=FeFp. The rotation and stretching of the lattice are taken into account through the elastic deformation gradient Fe [4].The plastic deformation gradient Fp includes the constant volume plastic deformation without disturbance of the crystal lattice.

If the convergence criteria in step (V) are fulfilled the state variables are updatedfixed and the next load step is calculated.(cid:7) ;7DE,[Results and Discussion Process Boundary ConditionsThe applied boundary conditions in UC are shown in Figures 2 & 3. The contactboundary conditions (Figure3) with coulomb friction can be found from the literature. [5]Figure2: Macroscopic Boundary conditionsFigure3: Virgin layer and substrate scenarioModel Results The model has been tested while predicting the joining behavior of Al 3003 H18 thin sheet with thickness dimension of 150 µm on a substrate of the same material. For all practical purposes, the thickness dimension of the substrate has been assumed to be 1mm [6] and then a representative volume (RVE) has been generated with dimensions of 2mmx1.15mmx2mm as shown in figure 4.

As shown in figure 4, the finite element mesh for deformation analysis has been made finer near the interface (~5µm length in the loading direction) since it has a very high amount of rigidity to dislocation motion due to high interfacial energy which will further lead to huge deformation gradients between the interface and the bulk material. Other input parameters were applied uniform normal pressure of 25 MPa [6], ultrasonic sinusoidal wave with Amplitude of 16 µm and frequency of 20 kHz. The deformation behavior has been analyzed for one cycle since the model is accurate to the order of dislocation content and their mesoscopic interactions and due to its computational cost. Material parameters for the constitutive model at the integration point level have been included in Table 1 [3]. Parameters which have been ignored in the current study are weld speed (v), partial contact between sonotrode and Al3003 H18 foil and between H18 foil and H18 substrate.

Parameters which have been ignored in the current study are weld speed (v), partial contact between sonotrode and Al3003 H18 foil and between H18 foil and H18 substrate. These parameters will soon be included for better accuracy in results.elementFigure4: FE Model for simulating the Ultrasonic consolidation process conditions for Al3003 H18 alloy. Red color denotes the foil and the blue color denotes the substrate. Substrate dimensions are 2mmx1mmx2mm and foil dimensions are 2mmx150µmx2mmThe objective of this study is to understand the interface characteristics between the substrate and the foil. The substrate-foil interface conditions have been modeled by computing the average frictional energy stored at the interface per unit of time, which equals 1.912x10-9 J/time step. Since the element size of the FE model near the interface is 103 times bigger than the actual interface length (~b=10-10m), the energy stored at the interface is scaled to 1.912x10-12 J/time step.

Since the element size of the FE model near the interface is 103 times bigger than the actual interface length (~b=10-10m), the energy stored at the interface is scaled to 1.912x10-12 J/time step. The component of this energy in the slip direction gets added to the activation energy barrier required for the dislocation slip, Qslip (~3x10-19 J/time step) for all the integration points sharing the interface.

This has been compared to a normal metallic interface for example a grain boundary (~1/2Gb3~6.65x10-19 J/time step) which is scaled (6.65x10-22 J/time step) and resolved in the slip direction and the deformation behavior has been observed.Material Parameter Qslip Qbulk c1c2 c3 c4 c5 c6 c7 c8 c9Physical MeaningEnergy barrier for slip Energy barrier for climb (activated at higher temperatures) Constant for passing stress (due to in-plane dislocations) (Eqn3) Constant for Jump width (Eqn4) Constant for obstacle width (Eqn4) Constant for lock forming rate (Eqn7) Constant for athermal annihilation rate (Eqn7) Constant for thermal annihilation rate (Eqn7) Constant of dipole forming rate (Eqn7) Constant for non-linear climb of edge dislocation (Eqn7) Constant for energy scaling at the interfacePrescribed Value3.0x10-19 J 2.4x10-19 J 0.12.0 1.0 1.5x107 m-1 10.0 1.0x10-30m-1 1.0x107m5sc8 0.3 10-3It has been seen with progressive deformation that the interface with frictional contact between the substrate and the foil experiences a bigger magnitude of the geometrically necessary dislocations in the edge dislocation normal direction by a factor of 100 times.From a physical viewpoint, a higher density of geometrically necessary dislocations in the edge normal dislocation direction means that each mobile dislocation of edge nature on its slip plane experiences a higher resistance to its motion leading to their immobile accumulation and resistance to slip plane rotation at the interface (figure 5).

Therefore, a higher geometrically necessary dislocation density along the frictional interface compared to the grain boundary interface leads to its grain fragmentation in the current and later cycles. The region of grain fragmentation could be identified as the thin regime in red color after 3/4th and 1 complete cycle of the deformation is complete (figure 6, right). The thickness of this regime is 2 mesh layers in the loading direction or 10µm which matches with the grain fragmentation regime in [1] Moreover, the yield and the stress-strain curves in both the cases are identical (~160MPa (initial), figure 7) because in the small strain regime, it is only a function of the statistically stored dislocation density which measures the multiplication of initially frozen dislocations which are trapped in the material in the as received condition.

Since the initial stored statistical density is ~1014/m2 and the maximum absolute change observed during progressive deformation is no larger than ~105/m2, therefore a change in the yield stress has not been seen when the frictional contact has been compared to the grain boundary interface.Figure5: Physical interpretation of dislocation content at the interfaceT=1/4 CycleT=1/2 CycleT=3/4 CycleT=1 cycle.Figure6: The edge normal dislocation evolution at the H18-H18 interface. Pictures at left indicate grain boundary interface and pictures at right indicate the friction boundary interface. Clearly, the rotation restricting dislocation content at the interface is 100 times stronger in the frictional case than in the grain boundary case. This further leads to grain fragmentation.Figure7: The stress-strain responses of the H18 substrate-H18 foil system with frictional and grain boundary interfaces in red and green respectively.

This further leads to grain fragmentation.Figure7: The stress-strain responses of the H18 substrate-H18 foil system with frictional and grain boundary interfaces in red and green respectively. Notice that in the initial regime, the behavior is primarily dictated by the statistically stored dislocation density.Conclusions and Future workIt can be concluded from the present work that the dislocation density based non-local non-linear finite element modeling can be used to analyze the Ultrasonic Consolidation process down to the mesoscopic level and many evolutionary variables, for example plastic deformation, dislocation content and grain fragmentation, can be predicted. The downside of this analysis is the time involved for computing the variables of interest close to interface because of finer mesh size. This problem can be eradicated by a robust homogenization strategy and its mathematical framework is already under development.

This problem can be eradicated by a robust homogenization strategy and its mathematical framework is already under development. A detailed TEM study will be conducted to identify the material parameters as shown in Table 1 when other material systems will be joined using the UC technique. The present model will be validated against the deformation response of these material systems and then converted to a tool for predicting the deformation response of new systems and the feasibility and amount of bonding which could be achieved using the UC technique in those cases.

Other regimes of research interest would be to look at better statistical mechanics based ensembles and phase field study to understand the nature of the dislocation locks produced on dislocation multiplication which further restricts their motion in the slip direction.Once these objectives are fulfilled, the combined experimentation and computationapproach towards understanding the UC technique could be used as a tool in1. Better understanding of the effect of UC process parameters on interfacial characteristics 2. Performing analytical optimization of process parameters for a given set of materials to be bonded; dramatically reduces the need for full factorial studies of process parameters.3. Ability to model the effects of changes in sonotrode design and 4. Predicting the suitability of new materials for UC processing.References1. Johnson, K. Interlaminar subgrain refinement in ultrasonic consolidation.

SLOCATION DENSITY CRYSTAL PLASTICITY BASED FINITE ELEMENT MODELING OF ULTRASONIC CONSOLIDATIOND. Pal*, B. E. Stucker+Department of Mechanical & Aerospace Engineering, Utah State University, Logan, UT 84322 +Department of Industrial Engineering, University of Louisville, Louisville, KY 40292AbstractA dislocation density based constitutive model has been developed and implemented into a crystal plasticity quasi-static finite element framework. This approach captures the statistical evolution of dislocation structures and grain fragmentation at the bonding interface when sufficient and necessary boundary conditions pertaining to the Ultrasonic Consolidation (UC) process are prescribed.The hardening is incorporated using statistically stored and geometrically necessary dislocation densities (SSDs and GNDs) which are dislocation analogs of isotropic and kinematic hardening respectively.

Since, the macroscopic boundary conditions during UC involves cyclic sinusoidal simple shear loading along with constant normal pressure, the cross slip mechanism has been included in the evolution equation for SSDs. The inclusion of cross slip promotes slip irreversibility, dislocation storage and, hence, cyclic hardening during the UC. The GND considers strain-gradient and thus renders the model size-dependent. The model is calibrated using experimental data from published refereed literature for simple shear deformation of single crystalline pure aluminum alloy and uniaxial tension of polycrystalline Aluminum 3003-H18 alloy. The model also considers the tension-compression asymmetry in case the model is applied for deformation processes in hexagonal close packed pure Titanium and its alloy counterparts which will be investigated further in our proposed research program.

One of the significant macroscopic contributions from this model development is to successfully accommodate the elasto-plastic contact problem involved in UC.The model also incorporates various local and global effects such as friction, thermal softening, acoustic softening, surface texture of the sonotrode and initial mating surfaces and presence of oxide-scale at the mating surfaces which further contribute significantly specifically to the grain substructure evolution at the interface and to the anisotropic bulk deformation away from the interface during UC in general. The model results have been predicted for Al-3003 H- 18 alloy undergoing UC. A good agreement between the experimental and simulated results has been observed for the evolution of linear weld density and anisotropic global strengths macroscopically.

A good agreement between the experimental and simulated results has been observed for the evolution of linear weld density and anisotropic global strengths macroscopically. Similarly, microscopic observations such as embrittlement due to grain substructure evolution and broken oxide layer at the UC interface has been also demonstrated by the simulation.IntroductionAs a direct result of ongoing research efforts in ultrasonic consolidation (UC) worldwide, it has become apparent that a new approach to modeling of UC bonding is needed.

A model which provides a better understanding of the effects of process parameter changes on grain refinement,plastic deformation and bonding during UC will better enable researchers to predict which materials will bond, how the mechanical properties of UC-produced parts can be improved, and how to better design the next generation of UC equipment.The continuum properties of parts made using UC are strongly dependent upon the micromechanics of the bonded interface [1]. Interfacial-scale microstructures can be studied fundamentally using electron microscopy and can be used to correlate atomic and mesoscopic mechanisms of deformation to their continuum counterparts.

Interfacial-scale microstructures can be studied fundamentally using electron microscopy and can be used to correlate atomic and mesoscopic mechanisms of deformation to their continuum counterparts. A dislocation density-based crystal plasticity finite element model (DDCP-FEM) can capture the statistical distribution of dislocations, partials and various deformation mechanisms at the bonding interface as inputs to predict macroscopic deformation and mechanical property profiles as a function of energy input characteristics. These input characteristics are a function of the process parameters used in a UC machine, namely vibration amplitude, normal force, ultrasonic frequency, welding speed, sonotrode geometry and temperature.Problem FormulationIt has been shown that material sheets subjected to UC undergo inhomogeneous plasticdeformation through their thickness [1]. Classical continuum plasticity theories do not fully explain this phenomenon [2].

Classical continuum plasticity theories do not fully explain this phenomenon [2]. Therefore, a study of strain localization and grain refinement at the material interfaces during UC bonding is required. The following steps lead to the calculation of these localized strains and their effects.1 Large Deformation Quasi-static formulationThe deformation map in space and time is described by the total deformation gradient tensor F (Figure 1). Applying the Kroner-Lee assumption, F is decomposed into elastic Fe and plastic gradient Fp tensors using multiplicative operator theory(Eqn 1)The plastic deformation gradient Fp includes constant volume plastic deformationwithout disturbance of the crystal lattice.

Elastic distortion and rigid rotation of the lattice are described by a unique intermediate configuration free of local stresses.2 The non-local dislocation density motivated material modelThe flow response for dislocation density motivated crystal plasticity modeling in a givenslip system 'α' is given by (Ma 2006):̇{̇(| |)+ ( )| | | |(Eqn 2)̇ is the upper limit of the shear rate for the case where thewhere the pre-exponential variable Boltzmann factor is equal to 1, which can be found using:̇√(Eqn 3)Figure1 Multiplicative decomposition of the total deformation gradient, F=FeFp.

The rotation and stretching of the lattice are taken into account through the elastic deformation gradient Fe [3].and the passing stress,, caused by parallel dislocations can be found using:√(Eqn 4), at 0K caused by forest dislocations can be found using:and the cutting stress,√(Eqn5)where Qslip is the effective activation energy for dislocation slip.The incompatibility in plastic deformation gradient and non-local geometrical non- which computes the geometrically necessary dislocationslinearity is introduced using required to maintain continuity throughout the material. The evolution law foris:̇‖ ( ̇) ̃ ‖(Eqn6)Material hardening at an integration point is both a function ofand are generally linear in shear(statistically stored dislocation density).

The evolution laws for rate (Eqn7).√ ̇̇̇̇ ()| |) ( ̇ ) (Eqn7)(The various constants used in this modeling can be attributed to physical phenomena asshown in Table 1.Material Parameter Physical MeaningPrescribed ValueQslipEnergy barrier for slip3.0x10-19JQbulkEnergy barrier for climb (activated at higher temperatures)2.4x10-19Jc1constant for passing stress (due to in- plane dislocations) (Equation 3)0.1c2constant for jump width (Equation 4) 2.0c3constant for obstacle width (Equation 4)1.0c4constant for lock forming rate (Equation 7)1.5x107m-1c5constant for athermal annihilation rate (Equation 7)10.0c6constant for dipole forming rate (Equation 7)1.0x10-30m-1c7constant for thermal annihilation rate (Equation 7)1x107m5c8c8constant for non-linear climb of edge dislocation (Equation 7)0.3c9constant for energy scaling at the interface10-3Table1 Physical Interpretation for various constants used in the Constitutive model3 Global and Local Solution StrategiesThe global and local solution strategy developed for solving the crystal plasticityconstitutive model in section 2.2 is shown in Table 2.Table2 Solution strategy for solving the crystal plasticity constitutive model with Nonlinear FEM.First, the applied loading is applied incrementally as load steps.

At the beginning of each load step (k+1), various nodal and integration point variables such as displacement ( ), plastic deformation gradient ( ), SSD and GND are initialized to their converged counterparts obtained in the previous load step. After initialization, an iterative procedure is used to solve for the local integration point variables. In every iteration (i), a local integration is performed via the B-Bar scheme to compute the stress, plastic deformation gradient and the hardness due to immobilized dislocation content (SSD). The stress is then resolved on a slip system to obtain the resolved shear stress (RSS, ) followed by plastic shear strain rate ̇ which further leads to the calculation of the stiffness matrix (C) at the integration point followed by a nodal updated tangent modulus (K) as a function of displacement. The tangent modulus is a measure of local stress with respect to local strain.

The tangent modulus is a measure of local stress with respect to local strain. This tangent modulus in uniaxial tensile plasticity is lower than the Young’s modulus (E) since the force required to glide dislocations is less than the force required for moving perfect atomic planes. After the calculation of nodal tangent modulus, the internal nodal force based on stresses at the integration point and the integration algorithm (B- Bar) is used to calculate the force residual (Q). Solving the matrix-inversion problem posed in B(II) in table 2, an increment in the displacement ( ) is obtained for the current iteration (i). The increment in displacement is added to the current displacement for updating its value for the next iteration (i+1), if required. The increment in GND is calculated at this stage. The procedure is illustrated in Figure 2. First, the plastic deformation gradient ( ) is linearly extrapolated from local integration points to the corresponding nodes.

The procedure is illustrated in Figure 2. First, the plastic deformation gradient ( ) is linearly extrapolated from local integration points to the corresponding nodes. Next, the nodal value of is averaged out ) is since the node under consideration is connected to other elements. The curl ( ̇ calculated and returned to the integration points from their corresponding nodes in step 3. The ) at integration points is multiplied by appropriate material variables to obtain curl ( ̇ GND for screw and edge dislocations in their normal and tangential directions respectively. To ensure that the local integration point has converged, the increments in , SSD and GND are compared against their respective threshold upper bounds. This is followed by global convergence, where the absolute value of the force residual (Q) is compared against a global threshold.

This is followed by global convergence, where the absolute value of the force residual (Q) is compared against a global threshold. Once both the local and global convergences are ensured, various nodal and integration point variables are predicted for the next load step; and the same sequence described here is repeated for the next load step (k+1).Figure2 Procedure to determine the slip-rate gradients within 3D trilinear elements with B-bar integration. The interpolation of the * ̇ + is performed via an eight-noded solid element with (2×2×2) integration.4 Process Boundary ConditionsThe model formulations and solution strategies in sections 2.1, 2.2 and 2.3 have been validated for simple boundary conditions such as simple shear and normal compressive loading [4].

The model results such as stress evolution, GND or inhomogeneous plastic deformation evolution and SSD or homogenous plastic deformation evolution are in good agreement with the experiments (Ma 2006) and (Siddiq 2007) respectively. UC process requires a simultaneous application of the abovementioned boundary conditions along with local interfacial contact based boundary conditions. Therefore based on the accuracy of the model established by validation of the results in the simple cases, more complex boundary conditions representing the UC process have been applied to the model and are discussed in this section.

The model predictions for these complex boundary conditions and their comparison with the experiments will be discussed in section 3.4.1 Global process boundary conditionsThe schematic of the UC process in Figure 3 illustrates the global boundary conditions, namely normal compressive force, oscillation amplitude, weld speed and initial conditions like foil microstructures and surface roughness at the interface.The schematic in Figure 3 has been converted into mathematical boundary conditions asshown in Figure 4.Figure3 Schematic of the UC processThe normal compression in UC is around 1200N -2000N for Aluminum alloy. The applied shear amplitude is of the order of 1µm-25 µm with foil thicknesses between 50 µm-200 µm (a typically used foil thickness is ~150 µm) and initial mating surface roughness varying from smooth (Ra~0.1 µm) to heavily surface damaged (Ra~5 µm).

The mentioned surface conditions are only observed in the top part of the bottom foil because it has been freshly deformed by the sonotrode during deposition of the previous layer. The time period for oneoscillation is about 50 microseconds (corresponding to a sonotrode frequency of 20kHz). Since the travel speed of the sonotrode horn varies from 10-50 mm/s, the material point under UC loading experiences a maximum of 3000 material deformation cycles. Foils are typically composed of polycrystalline microstructures with random orientation as shown in [1]. The initial SSD has been assumed to be 5x1013m-2 [4]. The microscopic parameters required for simulations can be extracted from Table 1. Since, orientation imaging microscopy (OIM) data for the foils were not available, a random cubic polycrystalline microstructure has been assumed.

Since, orientation imaging microscopy (OIM) data for the foils were not available, a random cubic polycrystalline microstructure has been assumed. The bottom foil and the top foil are comprised of 12 randomly generated crystals each, with the top foil having the same granular microstructure as the bottom foil.4.2 Interfacial contact boundary conditionsFor interfacial contact with friction, the contact boundary conditions can be formulated asfollows:( )(Eqn 8)where and are displacements of the nodes on the top of the bottom and bottom of the top foils in the vicinity of the interface respectively. denotes the normal vector perpendicular to the interface. The normal vector is in the y-direction for this case.Figure4 Schematic showing Ultrasonic Consolidation global boundary conditions. The bottom of the model is held rigid.

The normal vector is in the y-direction for this case.Figure4 Schematic showing Ultrasonic Consolidation global boundary conditions. The bottom of the model is held rigid. This assumption is only true for 2 foils on a large plate or a foil laid over a large plate.In addition to equation 8, the normal traction components on the bottom of the top foil and top of the bottom foil should be compressive and tensile respectively.

Also, from the point of view of force equilibrium, both of these traction components should be equal and opposite to each other as shown in equation 9.(Eqn 9)Equations 8 and 9 further lead to the Kuhn-Tucker conditions [11] for the normal tractionand normal displacement components as demonstrated in equation (10)( )( ) =0(Eqn 10)Further, similar to the normal traction components, the tangential components should bealso in equilibrium (equation 11).(Eqn 11)where is the tangential projection operator, given as:( )(Eqn 12)where denotes the cross product of two vectors.For friction idealized by the Coulomb law, the maximum frictional force supported by theinterface is:| |(Eqn 13)where is the coefficient of friction.

The equations to be satisfied for the tangential component of interfacial tractions and displacements are as follows:|| ||(Eqn 14)|| || ( )(Eqn 15)( ) || || ( )(Eqn 16)Results and Discussion1 Effects of Friction at the UC interfaceTwo cases have been modeled to understand the effects of friction at the UC interface, one with a pure stick condition and the other with an appreciable amount of friction, replicating the friction behavior when the constituent foils are in contact with each other. The initial mating surface roughness has been fixed at an average roughness, Ra~5 µm with foil thickness being 150 µm. The boundary conditions applied are (a) normal compression of 1800 N and (b) oscillatory shear amplitude of 16 µm. The initial GND has been assumed to be 0 throughout the microstructure since the grain boundary area is negligible compared to the entire volume. The GND evolution in the bulk for one complete deformation cycle has been shown in Figure 5.

The GND evolution in the bulk for one complete deformation cycle has been shown in Figure 5. Figure 5(a) illustrates the initial configuration for both the pure sticking and frictional sliding boundary conditions. During progressive loading in the pure sticking case, the mating interface experiences no resistance to the shear deformation. This leads to equal tangential displacements of the top and bottom foils in the vicinity of the interface. Therefore, the GND starts developingat the top surface of the top foil. On the contrary, in the frictional sliding case, the bottom foil in the vicinity of the interface experiences significant resistance due to friction. This further leads to much smaller tangential displacement of the bottom foil compared to the top foil in the vicinity of the interface. Henceforth, the GND is higher at the interface compared to the top surface of the top foil in this scenario.

Henceforth, the GND is higher at the interface compared to the top surface of the top foil in this scenario. The GND evolution in both pure sticking and frictional sliding scenarios has been demonstrated in figures 5(b) through (e) corresponding to (1/4)th, (1/2), (3/4)th and 1 full deformation cycle during UC. It is also observed that the maximum GND after one complete cycle (figure 5(e)) in the frictional sliding case is greater than the pure sticking case by a factor of 2 suggesting more inhomogeneous deformation in the former. The motivation behind demonstrating these particular instants (figures 5(a)-(e)) comes from the fact that the applied shear loading either changes its direction or its sign at these instants. Also, the extent of GND evolution in the frictional sliding case limits itself to about 20 µm below the top surface of the bottom foil in the normal loading direction (y-axis).

Also, the extent of GND evolution in the frictional sliding case limits itself to about 20 µm below the top surface of the bottom foil in the normal loading direction (y-axis). Therefore, this region experiences the maximum the inhomogeneous plastic deformation, and kinematic hardening inhomogeneous plastic and kinematic hardening is directly proportional to the √ (Ma 2006). Moreover, continued GND evolution at the interface means a higher production of dislocations with either a positive or a negative sign. Therefore, a very high number of these GNDs lead to their annihilation and formation of new relaxed equiaxed subgrains. This phenomenon is termed dynamic recrystallization [6].

Therefore, a very high number of these GNDs lead to their annihilation and formation of new relaxed equiaxed subgrains. This phenomenon is termed dynamic recrystallization [6]. The emergence of recrystallized subgrains can be confirmed from experiments [1] where it has been observed that significant grain fragmentation has taken place near the interface on the top part of the bottom foil with a good amount of plastic flow beneath the region, as shown in figure 6.sinceThis continued evolution of GND at the interface in the friction assisted case results in grain fragmentation (subgrain formation) by the end of 3000 material deformation cycles, since the average GND in the 20 µm strip below the top surface of the bottom foil reaches a value ~ 6x1015 dislocations/m2 in the top part of the bottom foil (Figure 7).

The sub-grain diameter corresponding to this dislocation density is ~ 2.28 µm [4].Figure 5(a) GND evolution at T=0 (i) Pure sticking (ii) Frictional sliding385Figure 5(b) GND evolution at T=1/4th cycle (i) Pure sticking (ii) Frictional sliding386Figure 5(c) GND evolution at T=1/2 cycle (i) Pure sticking (ii) Frictional sliding387Figure 5(d) GND evolution at T=3/4th cycle (i) Pure sticking (ii) Frictional sliding388Figure 5(e) GND evolution at T=1complete cycle (i) Pure sticking (ii) Frictional sliding389Figure6 Ion beam induced secondary electron micrograph of DBFIB-etch result for UC interface showing extent and depth of sub-grain size and morphology change as a result of the plastic deformation [1]2 LWD evolution with cyclic deformationLinear Weld Density measures the amount of bonded length to the total interfacial length, which is inversely proportional to the average gap between the mating surfaces at the UC interface.

LWD generally increases with the amount of cyclic deformation. If the gap becomes zero, both the surfaces adhere and start plastically deforming together. Since the applied shear loading is cyclic during UC processing, the interface may lead to a fatigue type of failure and dynamic recrystallization at the interface when subjected to higher normal loads. Although the gap closure phenomenon has been included in the model, the prolonged fatigue phenomenon has not been included in the current set of simulations and is left for future work.To quantify the effects of processing parameters such as oscillation amplitude and normal compressive load on LWD evolution, a design of experiments (DOE) approach was used by [5] to systematically evaluate the effects of process parameters and to identify the optimum parameter combination. Specific levels for each of the parameters were selected based on preliminary experiments, machine setting limits, and available published information.

Specific levels for each of the parameters were selected based on preliminary experiments, machine setting limits, and available published information. A Taguchi L16 orthogonal array was utilized in his study to determine the effects of individual process parameters.

Welding occurred along the 100 mm direction. Metallographic sample locations are shown in the picture [5].2.1 Effect of oscillation amplitudeThe L16 Taguchi experiments conducted by [5] are an ideal DDCP-FEM model validation tool for UC. To investigate the effect of oscillation amplitude on LWD, four distinct simulations have been carried out with an applied normal load of 1750N for 3000 deformation cycles and applied ultrasonic amplitudes of 10, 13, 16 and 19 µm. The foil thickness was assumed to be 150 µm, with an average roughness of Ra~5 µm for the top part of the bottom foil. Figure 9 shows the average gap evolution between the mating surfaces as a function of increasing amount of cyclic deformation. It can be seen that for most amplitudes, the average gap no longer decreases after approximately the 10th cycle. For smaller amplitudes, the average gap decreases with increasing amplitude up to 16 µm.

For smaller amplitudes, the average gap decreases with increasing amplitude up to 16 µm. Above 16 µm the gap increases with increasing amplitude.LWD has not been explicitly predicted in the simulation since the model treats gaps between the top and the bottom foils as 2 dimensional quantities whereas LWD is a one dimensional, linear quantity. However, the average gap can be used as an effective parameter for comparison with LWD, as LWD and average gap should show all of the same trends. A quantitative calibration between the average gap and LWD is left for future work.As can be seen from figure 10, the average gap shows the same trends shown by [5], first increasing and then decreasing as a function of the applied ultrasonic amplitude. The optimum amplitude where the average gap is found to be the least was 16 µm for the applied boundary conditions.

The optimum amplitude where the average gap is found to be the least was 16 µm for the applied boundary conditions. This is in good agreement with the experiments since the peak LWD has been observed at the same oscillation amplitude of 16 µm.Figure9 Gap evolution (in µm) against number of deformation cycles with varying ultrasonic shear amplitudes.Figure 10: Effect of processing parameters on linear weld density, LWD vs. oscillation amplitude from [5]2.2 Effect of Normal forceSimilarly, to understand the effects of increasing normal load on gap closure/LWD, four applied normal loading scenarios have been considered. The applied normal loads were 1450, 1600, 1750 and 1900 N. The applied ultrasonic amplitude was 16 µm. The foil thickness and average roughness on the top part of the bottom foil were kept at 150 µm and Ra~5 µm respectively.

The applied ultrasonic amplitude was 16 µm. The foil thickness and average roughness on the top part of the bottom foil were kept at 150 µm and Ra~5 µm respectively. Again, it can be seen (Figure 11) that there is no appreciable change/decrease in the average gap as a function of increasing cyclic deformation past the end of the 10th cycle for most cases.It has been observed by Janaki Ram et al. [5] that when increasing the applied normal force, LWD increases up to a load of 1750 N. A further increase in normal load, however, lowered LWD. The most probable reason behind this LWD decrease is increase in the inhomogeneous plastic deformation with load near the top surface of the bottom foil. This leads to very high amounts of dynamic recrystallization which leads to dynamic failure of the bonded region at the interface.

Therefore, in order to predict the experimental results, it will become necessary to incorporate fatigue crack nucleation and growth phenomena within the current DDCP-FEM.The inclusion of fatigue phenomena is non-trvial for the following two reasons:The model assumes that once two nodes from the top and bottom surfaces are in contact with each other, the surfaces will be metallurgically bonded.A traction-separation law to mimic crack propagation with increasing cyclic deformation is not available from the literature.In future work, a traction separation law will be formulated as a function of involved fatigue damage and further crack propagation to try to capture the bond degradation seen during experiments at higher loads.In conclusion, the effects of oscillation amplitude and normal force on LWD evolution have been correctly modeled in this section with the exception of decreasing LWD with increasing normal force beyond a certain magnitude.

Though the simulation parameter (average gap) characterizing the bonded region has not been calibrated quantitatively with LWD, it provides a good heuristic for bond characterization since the interfacial void closure is a 2 dimensional area based phenomenon.Figure11 Gap evolution (in µm) against the number of deformation cycles with varying normal loading scenarios.Figure12 Effect of processing parameters on linear weld density, LWD vs. Normal force from [5]3 Effects of Mating Surface RoughnessThe initial mating surface roughness at the top surface of the bottom foil is another important parameter for determining the amount and type of plastic deformation at the mating interface. The reason behind a rough top surface of the bottom foil during ultrasonic consolidation is because of the plastic work previously done by the sonotrode [8].

The reason behind a rough top surface of the bottom foil during ultrasonic consolidation is because of the plastic work previously done by the sonotrode [8]. Two distinct simulations have been carried out to investigate the effects of surface roughness.The simulation boundary conditions are: (a) normal compressive load of 1750N, (b) ultrasonic simple shear amplitude of 16 µm, (c) foil thickness of 150 µm, and (d) microstructure as shown in Figure 4. The top surface of the bottom foil has been assumed to have a roughness of Ra~0.1 µm in one case and Ra~5 µm in the second case. The lower roughness scenario represents two cases: (a) virgin foil depositing over a machined base plate, and (b) virgin foil deposition over a previously deposited and subsequently machined foil.

The lower roughness scenario represents two cases: (a) virgin foil depositing over a machined base plate, and (b) virgin foil deposition over a previously deposited and subsequently machined foil. It has been observed that the foil with lower surface roughness closes instantaneously in the first scenario whereas it takes around 10 deformation cycles to close the initial mating surface gap in the second scenario, as shown in figure 12. This further indicates a higher LWD in the low roughness case compared to the higher roughness case. This result is in good agreement with experiments [5], where the LWD is higher in between the machined base plate and virgin foil and lower in between the two layers comprising of sonotrode driven surface damaged foils.

Similarly, GND evolution in the 20 µm strip below the interface sees a much higher rate of evolution in a higher roughness scenario than a lower roughness scenario as shown in figure 13, suggesting a greater amount of subgrain formation near the top surface of the bottom foil in the former case than in the latter case. This observation is in good agreement with experiments performed by [5] where it has been observed that the top surface of the bottom foil undergoes minimal subgrain formation since a smoother interface generally provides much less frictional resistance and in turn a lower value of the lattice curvature.

The average subgrain diameter in the lower roughness scenario is ~4.29 µm which is about 2X more compared to the higher roughness scenario ~2.28 µm, establishing a close match with the experimental observations [12].Another important aspect is the mating interface shape, which has a higher rate of curvature in the higher roughness scenario than the lower roughness scenario, as shown in figure 14, after 3000 cycles. Again, this result can be validated using [5] depicting the interface shapes at the end of around 3000 cycles with lower and higher initial average roughness, Ra ~ 0.1 and 5 µm respectively. Clearly, the mating interface in the high surface roughness scenario shows a very high curvature and provides a good representation of the section at the interface demarcated in figure 13(a).

Clearly, the mating interface in the high surface roughness scenario shows a very high curvature and provides a good representation of the section at the interface demarcated in figure 13(a). Similarly, the planar interface in a low roughness scenario is a representative section of the interface demarcated in figure 13(b).Figure 11: Gap evolution (in µm) against number of deformation cycles with varying bottom foil roughnessFigure12 GND evolution in the 20 µm strip below the interface against number of deformation cycles with varying bottom foil roughness. A higher GND at this location means more prominent subgrain formation.

A higher GND at this location means more prominent subgrain formation. This result is in agreement with the experimental observations demonstrated in [5].Figure13 Interface evolution against number of deformation cycles with varying bottom foil roughness.4 Evidence of closure and relaxation at the interfaceIt has been observed that the gap between the top and the bottom foil being consolidated by UC undergoes closure and relaxation [7] based on its location at the interface and local stress state. Since, the gap between the top and bottom foil is prominent in the first few cycles, the closure-relaxation effect is also more pronounced in the earlier cycles.The mechanism of gap closure and relaxation as a function of location and state of stress is summarized in figure 14. In the first quarter cycle at the rightmost locations, the top foil just above the interface undergoes a compressive state of stress and the bottom foil goes through a tensile state of stress.

In the first quarter cycle at the rightmost locations, the top foil just above the interface undergoes a compressive state of stress and the bottom foil goes through a tensile state of stress. Also, at this point in time, the top foil at the leftmost locations just above the interface see a tensile state of stress and the bottom foil sees compressive stresses. Henceforth, the gap closes at an appreciable rate for the rightmost locations of the interface whereas it remains the same or slightly increases for the leftmost locations. In between the first quarter and half cycle, the top foil at the rightmost locations just above the interface develops a tensile state of stress whereas the bottom sees a compressive state of stress. Hence, the gap at these locations increase whereas the gap at the leftmost locations close because the top foil just above the interface develops a compressive state of stress and the bottom experiences tensilestresses.

The gaps at the leftmost locations continue to close because of the continued state of stresses in the top and bottom foils in the vicinity of the interface for the time between one half cycle and the last quarter of the complete cycle. Meanwhile at the rightmost locations, the top foil again starts seeing a compressive state of stress and the bottom foil experiences a tensile state of stress leading to gap closure. This state of stress continues for the rightmost locations in the last quarter of the complete cycle leading to further closure of the gap at these times. On the contrary, in the last quarter of the complete cycle, the leftmost locations again start seeing a tensile state of stress in the top foil and compressive stresses in the bottom foil leading to gap increase or gap relaxation.Figure14 The gap closure and relaxation as a function of location and applied loading at the mating interface. The arrows denote the direction of applied simple shear in a deformation cycle.

The arrows denote the direction of applied simple shear in a deformation cycle. The schematic represents the first cycle of UC deformation since the closure/relaxation mechanism is most pronounced in the transient stages.It is essential to take gap closure and relaxation mechanisms into account while the UC machine is in operation since it aids in deciding the optimum amplitude required for proper deposition of virgin foils over previously deposited foils/machined substrate. If the sonotrode horn is operated at lower amplitudes than the transverse locations at which gap closure occurs, the interfacial contact rate is lower than the optimum leading to a lower rate of average gap minimization and lower LWD. Similarly for very high amplitudes, the transverse locations at which gap relaxation happens, the interfacial distance between the foils increase at a faster rate than the optimum, again leading to a lower rate of average rate of gap minimization and lower LWD.

Henceforth, the optimum amplitude for average rate of gap minimization and LWD is achieved at an amplitude of 16 µm as above.In future work, the weld speed will be explicitly applied as a boundary condition with applied normal compressive load and oscillation amplitude. It is anticipated that at the longitudinal locations ahead of the sonotrode, the top foil in the vicinity of the interface will experience a compressive state of stress whereas the bottom foil will experience a tensile state of stress causing the longitudinal interfacial gap to close. On the contrary, at the longitudinal locations behind the sonotrode, the top foil at the vicinity of the interface will experience a tensile state of stress whereas the bottom foil will experience a compressive state of stress causing the longitudinal gap to either remain closed (since the gap has already been closed due to transverse oscillatory motion of the sonotrode) or to rip open if a high enough weld speed is applied.

Henceforth, a higher weld speed should decrease the LWD as the gaps behind the sonotrode starts to rip open. Another reason behind decreasing LWD with increasing welding speed is that the exposure time of the transverse weld cross-section to the sonotrode decreases with a higher welding speed, leading to incomplete bonding. The combined effect of gap closure and relaxation in the longitudinal direction along with incomplete bonding time on increasing welding speed has already been investigated and established by [5], and shown in figure 15. It is clear from figure 15 that the LWD decreases with increasing welding speed.Figure15 Effect of processing parameters on linear weld density, LWD vs. weld speed from [5].Conclusions and future work1 ConclusionsA DDCP-FEM model has been formulated and tested based on UC boundary conditions comprised of simultaneous simple shear and normal compressive force applied to polycrystalline Al 3003 alloy at ultrasonic frequencies.

The model was validated as follows:A higher normal force causes better bonding at the mating interface; though beyond a certain magnitude it results in a very high amount of in-homogeneous plastic deformation at the top surface of the bottom foil causing severe grain fragmentation and dynamicfatigue failure. This is reflected in LWD evolution, where LWD first increases with increase in normal force and decrease beyond a certain magnitude [5].A higher amplitude enhances better bonding at the interface. Although beyond a certain magnitude, it enhances the rate of gap relaxation, which overcomes the rate of gap closure leading to lack of fusion at the interface. Gap closure and gap relaxation are directly and inversely proportional to LWD evolution respectively. Therefore, it has been predicted that the LWD would first increase with the oscillation amplitude and thereafter decrease beyond a certain magnitude.

Therefore, it has been predicted that the LWD would first increase with the oscillation amplitude and thereafter decrease beyond a certain magnitude. This prediction is consistent with experimental results [5].A higher welding speed is beneficial from the point of view of quick deposition of virgin foils on top of a previously deposited foil(s)/machined base plate. Although with increasing welding speed, the rate of gap relaxation at longitudinal locations behind the sonotrode becomes significant leading these closed gaps to rip open. Also, a higher welding speed leads to less exposure time of the transverse weld cross-section to the sonotrode. This combined effect further results in lack of fusion at the interface and a lower LWD as manifested in [5].2 Future WorkThe current model is able to predict the deformation response during UC processing but requires additional features to be realized as an efficient tool for material and parameter optimization in UC.

The future accomplishments required in this paradigm are:To incorporate welding speed as a 3rd simultaneous applied load alongwith normal force and oscillation amplitude.To incorporate more number of foils being deposited on top of each other and their sequential consolidation. While doing this, the virgin foil for deposition will be assumed to have minimal initial GND and standard initial SSD content whereas for the previously deposited foils all the mesoscopic deformation variables such as SSD and GND will be stored from the previous deposition instant and applied as an initial condition for the fresh deposition.To formulate a traction-separation law which can quantify the delamination of foils as a function of applied normal and simple shear loading.

This is important from the point of view of optimizing the weld speed required to avoid crack propagation due to fatigue after bond formation.To perform Orientation Imaging Microscopy (OIM) of the as-obtained and post- consolidated UC-foils to generate realistic microstructures required for realistic simulations.To formulate a homogenized model based on current DDCP-FEM to predict deformation during processing and in-service conditions.

This is important since it will help with optimizing the initial orientation for the product build.To use the homogenized crystal plasticity model and non-destructive engineering metrics for closed-loop control (feed-forward and feed-back) of the UC machine.In its current form, the model is still under development and has to be validated for other crystalline types such as single and polycrystalline variants of body centered cubic (BCC) materials such as Molybdenum and Tantalum, hexagonal close packed (HCP) materials such as Titanium and mixed alloys containing both HCP and BCC counterparts such as Ti6Al4V. Thesealloys have been already demonstrated as potential materials for UC practice, though the biggest challenges to model these existing and new alloys which may serve as potential materials for UC are as follows:Determination of the active slip systems.

For HCP and BCC, because of their limited slip system activity, the active slip systems are a function of the Schmid factor (projection of loading axis on the slip system) and temperature. Therefore, the active slip systems can either be obtained from experiments and theoretical explanations provided in the literature, phase field modeling or insitu TEM experiments in which the resolved shear strain evolution can be monitored. Incorporation of mesoscopic deformation mechanisms, for example Nickel based superalloys such as Inconel 718, which demonstrate significant twin activity [9] and BCC materials exhibiting Peierls resistance to mobile dislocation motion [10].Determination of mesoscopic material parameters (Q’s and c’s shown in Table 1).  UC fabricated simple parts to validate model predictions.REFERENCES1. Johnson, K. 2008. Interlaminar subgrain refinement in ultrasonic consolidation.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDISSIMILAR VACUUM BRAZING OF ADDITIVE MANUFACTURED 17-4PH STAINLESS STEEL TO CONVENTIONAL 304 AND 17-4 STAINLESS STEEL BY BNI-2 BRAZING FILLER MATERIAL T.Göynük*,***, O.U. Onem*, E.Yasa**, İ. Karakaya ***ROKETSAN Missiles INC., P.K. 30 Elmadağ 06780 Ankara – Turkey ** University of Sheffield Advanced Manufacturing Research Centre, Factory of the Future Advanced Manufacturing Park Wallis Way, Catcliffe Rotherham S60 5TZ ***Middle East Technical University, Üniversiteler, Dumlupınar Bulvarı No:1, 06800 Çankaya/Ankara AbstractVacuum brazing process of 17-4PH stainless steel, which was manufactured using additive manufacturing, as well as conventional AISI304 and 17-4PH stainless steel alloys was investigated. The brazing process was conducted at 1050°C for 20 minutes under 10-6 Torr, using BNi-2 filler material.

The brazing process was conducted at 1050°C for 20 minutes under 10-6 Torr, using BNi-2 filler material. Various aspects of the vacuum brazed parts were analyzed, including their microstructure, wetting behavior and mechanical strength. To evaluate the wetting behavior, the contact angle and wetted area were measured using optical microscopy. The microstructures were examined using Scanning Electron Microscopy and Energy Dispersive Spectroscopy techniques. Additionally, tensile testing was performed on the joints to assess the influence of surface roughness and brazing of different materials on the strength of the brazed parts.IntroductionStainless steel alloys are of utmost importance in various industries because of their exceptional mechanical properties, corrosion resistance, and versatility [1].

Two widely recognized alloys within this category are 17-4PH stainless steel and AISI 304 stainless steel, which are highly regarded for their superior performance and diverse applications in aerospace. 17-4PH stainless steel is a precipitation-hardening alloy known for its remarkable strength, excellent corrosion resistance, and favorable mechanical properties even at elevated temperatures [2]. On the other hand, AISI 304 stainless steel belongs to the austenitic family, offering high corrosion resistance, good formability, and excellent weldability [3].In recent times, additive manufacturing has emerged as a disruptive technology that has revolutionized the manufacturing industry. This innovative manufacturing technique enables the creation of intricate designs and customized components while minimizing material waste and reducing lead times.

This innovative manufacturing technique enables the creation of intricate designs and customized components while minimizing material waste and reducing lead times. It provides unparalleled freedom in design and opens up new possibilities for optimizing the performance of stainless-steel alloys [4,5].Additionally, vacuum brazing has gained significant importance in the process of joining stainless steel components, particularly in environments with high temperatures and vacuum conditions [6]. Vacuum brazing ensures precise and clean bonding by eliminating oxides and impurities, resulting in strong and hermetic joints. It allows for the assembly of complex structures, reduces distortion, and improves the mechanical properties of the brazed parts. When performing vacuum brazing, the choice of filler material is critical to achieve reliable and successful joints.

When performing vacuum brazing, the choice of filler material is critical to achieve reliable and successful joints. BNi-2, a nickel-based filler material, has become prominent due to its advantageous characteristics, including a high melting point, excellent wetting properties, and compatibility with stainless steel alloys. The use of BNi-2 filler material facilitates the formation of robust metallurgical bonds, ensuring the integrity and durability of the brazed joints [7,8].Given the significance of these materials and processes, this study aims to investigate the vacuum brazing process of additive manufactured 17-4PH stainless steel, conventional AISI 304 stainless steel, and 17-4PH stainless steel alloys. The examination will focus on various aspects, including microstructure analysis, wetting behavior, permeability, pressure strength, and shear strength of the brazed parts.

The examination will focus on various aspects, including microstructure analysis, wetting behavior, permeability, pressure strength, and shear strength of the brazed parts. Through a comprehensive evaluation, this research intends to shed light on the influence of surface roughness, brazing techniques, and the effect of filler material on the performance of brazed components.350Materials and MethodsIn this study 20*40*3 mm test coupons were used for wetting and microstructure analysis. Additive manufactured 17-4PH steel test coupons were produced by pulsed mode selective laser melting method. Renishaw AM400 Selective Laser Melting (SLM) machine, which has a maximum size capacity of 250 × 250 × 300mm was used. This SLM equipment utilizes a Yb-Fiber pulsed-wave laser with a wavelength of 1064 nm and a maximum power of 400 W. The experiments were conducted in an argon environment. To support the structures, a 15 mm thick baseplate made of AISI 1040 steel was used.

The experiments were conducted in an argon environment. To support the structures, a 15 mm thick baseplate made of AISI 1040 steel was used. The production parameters of 17-4 materials manufactured by SLM are shared in Table 1.Table 1 Process parameters for 17-4PH stainless steel by Renishaw.Layer Thickness(μm) 30Power (W)200Hatching (μm)110Base Plate Temperature (°C) 20The chemical compositions, surface roughness values, and mechanical properties of all the materials used in this study are provided in the Table 2.

The profile can be divided into five distinct steps. Initially, the furnace was gradually heated to approximately 200°C and maintained at that temperature for approximately 15 minutes. This step aimed to eliminate any water-based contaminants and the filler material. In the second step, the temperature humidity from the system without causing any damage to351was raised to 540°C and held for ten minutes to remove organic substances from the surface. The third step involved heating the furnace to 950°C to ensure a uniform temperature distribution across the sample. At the fourth step, the filler material was melted. The final step involved cooling the furnace under a vacuum environment.Figure 1 Heating cycle of vacuum brazing process for BNi-2.Wetting analysis was done by optical microscope, while Scanning Electron Microscopy (SEM) together with Energy Dispersive Spectrometry (EDS) was used to evaluate microstructure analysis.

The geometry of the butt joint specimen used for mechanical tests and the geometry of the fixture used to align these specimens are shared in Figure 2.Figure 2 Butt joint and fixture geometry.Results and DiscussionWetting AnalysisWetting, an important criterion in evaluating the quality of brazing, has been examined by observing the spreading behavior of the same amount of BNi-2 paste on AM174, C174, and C304 coupons. The flow characteristics of brazing are influenced by several factors, including interalloying between the braze material and the substrate, diffusion of the braze material, dissolution of grain boundaries, and the formation of intermetallic compounds. These effects can be reduced by careful selection of the braze material, lowering the temperature, shortening the brazing duration, and increasing the cooling rate.

These effects can be reduced by careful selection of the braze material, lowering the temperature, shortening the brazing duration, and increasing the cooling rate. Conversely, they can be intensified by choosing a braze material with a high solubility limit in the substrate material or by raising the braze temperature and extending the brazing time [9].Braze wetting was assessed by measuring the contact angle and wetted area using optical microscopy. Surface flashing, blushing and wetting, were observed by visual inspection as shown in Figure 3. Braze alloys352that exhibit good wetting characteristics spread across the surface with a sharp contact angle at the edge of the molten pool, whereas braze alloys that do not wet tend to form spherical shapes on the surface. Ideally, when a braze wets the surface perfectly, it flashes the surface and forms a thin film with a contact angle close to zero.

Ideally, when a braze wets the surface perfectly, it flashes the surface and forms a thin film with a contact angle close to zero. Blushing, characterized by the separation of braze components, occurs due to a process called liquation, where elements with lower melting temperatures in non-eutectic alloys melt and separate from the braze. These separated elements flow across the surface, leaving behind the higher melting temperature elements. The occurrence of liquation observed in brazing filler material is dependent on whether the alloy is eutectic or not. Despite using the same brazing filler material, it exhibits different spreading characteristics on different materials in this study.

Despite using the same brazing filler material, it exhibits different spreading characteristics on different materials in this study. This difference can be attributed to various factors such as the metallurgical bond formed with the surface, the presence of contaminants or oxide layers on the surface, surface roughness, and other similar factors.Figure 3 Wetting characteristics of BNi-2 brazing filler on different materials.The analysis of braze wetting on AM174, C174 and C304 conducted using a parameter called the Wetting Index (WI). The WI is determined by multiplying the coupon area covered by braze flow with the cosine of the contact angle, as shown in equation (1), where A is the wetting area and θ is denoted for wetting angle. It is an empirical value that depends on the amount of braze material used in wetting tests and serves as a useful metric for comparing the wettability of brazes under identical test conditions.

It is an empirical value that depends on the amount of braze material used in wetting tests and serves as a useful metric for comparing the wettability of brazes under identical test conditions. All the results were tabulated in Table 4 including surface roughness, wetting index and wetting characteristics (melting, flashing, blush).

All the results were tabulated in Table 4 including surface roughness, wetting index and wetting characteristics (melting, flashing, blush). While wetting angle was measured during optical microscope, wetting area is calculated by using Image J program.𝑊𝐼 = {0, 𝜃 ≥ 90° 𝐴 × cos 𝜃 , 𝜃 < 90°(1)Table 4 Wetting analysis of BNi-2 filler material on AM174, C174 and C304.Base MetalSurface Roughness, Ra (μm)Contact Angle, θ (deg)Wetting Area (mm2)Wetting Index, Area*Cos (θ)WettingFlashing Blush2 - iN BAM1745.80.82975.2975.10✓✓C1741.02.55313.5313.19✓C3040.61.56396.0395.85✓✓Due to the nature of the additive manufacturing process, the surface roughness on parts is generally higher compared to conventional materials, despite efforts to reduce it through various process parameters. This difference in surface roughness has been observed to affect the wetting characteristics of brazing filler materials.

This difference in surface roughness has been observed to affect the wetting characteristics of brazing filler materials. Excluding conventional 17-4 material from the discussion, it can be observed that as surface roughness increases,353the wetting index also increases. However, it is also considered that at a certain point, excessive surface roughness may have a detrimental effect on the wetting behavior. Generally, roughening a flat surface was found to decrease the ability of a liquid drop to spread and wet the surface, unless the liquid had high inherent wettability, or the surface texture was extremely rough. Because surface roughness created energy barriers that a liquid had to overcome to wet the surface, and wetting was more likely to occur when these barriers were small compared to the enthalpy of the liquid. Surface roughness amplifies the influence of surface energy at the interface between the liquid, atmosphere, and the surface.

Surface roughness amplifies the influence of surface energy at the interface between the liquid, atmosphere, and the surface. A rough surface has a higher ratio of actual surface area to apparent surface area, which can enhance wetting when it is energetically favorable, or decrease wetting when it is energetically unfavorable [10]. Even though all materials have undergone the same cleaning procedure, it is considered that the surface of conventional 17-4 PH stainless steel may still retain an oxide layer from the solution treatment process, which is not completely removed. Therefore, it is evaluated that despite having higher surface roughness than AISI 304, the contact angle is higher, and the wetting index is lower for conventional 17-4 PH stainless steel. The ability of a braze material to flow should not be the sole criterion when selecting a braze alloy because some combinations of base metals and braze alloys do not promote free flow along the surface.

When the braze material and base metal undergo interalloying, it can increase the melting temperature of the filler above the brazing temperature, causing the alloy to solidify and impede further flow. Microstructure Analysis Brazing is a high-temperature process, when this joining technique is applied between two materials, the microstructure, diffusion, and interalloying between parent metal and filler metal should be thoroughly examined. To examine these properties the following brazing couples have been formed.C174 • C174 • C174 AM174 C174 • C304 • AM174 AM174 • C304 AM174 C304 • C304The microstructures of the brazed joints were analyzed using optical and SEM micrographs, with EDX mapping used to assess the elemental composition and measure the segregation of braze elements, agglomeration of precipitates, as well as the diffusion of brazing constituents within the base metals.

The optical microscope images and SEM micrographs of these brazing couples have been shared in Figure 4. Optical micrographs showing a clear contrast between the braze area and the base metal regions. In the optical microscope images, it was observed that darker-colored intermetallics formed within the brazing paste. These intermetallics are also visible in SEM images, exhibiting different distributions and morphologies across different brazing couples. It was determined that the diffusion and the resulting intermetallic formation within the base metal were more prominent in AISI 304 compared to 17-4 PH stainless steel. This enhanced diffusion facilitated stronger metallurgical bonding between the brazing material and the base metal. The diffusion depth in AISI 304 reached up to 30 microns, while it was measured at approximately 15 microns in 17-4 PH material.

The diffusion depth in AISI 304 reached up to 30 microns, while it was measured at approximately 15 microns in 17-4 PH material. Both conventionally and additively manufactured 17-4 PH materials exhibited similar behavior in terms of diffusion and microstructure. However, unlike AISI 304, voids were observed at the interface between the additively manufactured 17-4 PH material and the brazing material, which was attributed to surface roughness.354Figure 4 Optical and SEM micrographs of brazing couples.355As part of the elemental analysis, all brazing couples were analyzed through EDS mapping, and the results are presented below. In all joint interfaces, the formation of Cr and Mn carbides or borides within the brazing material was observed. These carbides and/or borides appeared larger and more angular in the presence of AISI 304 material in the brazing couple, whereas they exhibited a more homogeneous distribution throughout the joint region when 17-4 PH material was present.

In AISI 304-conventional or additive manufactured 17-4 PH joints, it was observed that the formed carbides generally concentrated towards the AISI 304 material, which had a higher carbon content.Figure 5 EDS mapping of C304-C304 brazing pair.Figure 6 EDS mapping of C174-AM174 brazing pair.356Figure 7 EDS mapping of C174-C174 brazing pair.Figure 8 EDS mapping of AM174-C304 brazing pair.357Figure 9 EDS mapping of AM174-AM174 brazing pair.Figure 10 EDS mapping of C304-C174 brazing pair.Mechanical TestsThe coupons prepared for mechanical testing were joined with BNi-2 brazing paste, following the brazing couples specified in the microstructure analysis. The use of a fixture ensured a fixed tolerance of 50-100 microns between the butt joint parts. All butt joint assemblies were tested at a speed of 1 mm/minute. The tensile stress- strain graph, derived from the raw data of the mechanical tests for the brazing couples, is provided in Figure 11.

All butt joint assemblies were tested at a speed of 1 mm/minute. The tensile stress- strain graph, derived from the raw data of the mechanical tests for the brazing couples, is provided in Figure 11. Additionally, for easier comparison, the complete data graph of the brazing couples is provided in Table 5.All test specimens ruptured at the brazing interface. It was observed that the elongation was higher in the brazing couples using AISI 304 stainless steel. Particularly, in the 17-4 PH materials produced through additive manufacturing, the elongation was found to be significantly low. It was observed that the gaps at the brazing interface, as seen in the microstructure analysis, contributed to this result.

It was observed that the gaps at the brazing interface, as seen in the microstructure analysis, contributed to this result. In the brazing couples using 17-4PH358material, the gaps in the interface acted as crack initiation points, resulting in both lower tensile stress and less elongation.In brazing couples using AISI304 material, higher strength and elongation values were achieved, along with the observed diffusion and metallurgical bonding. According to the mechanical test results, when comparing additive manufactured 17-4 PH material and conventional 17-4PH material, it was observed that the assemblies with conventional 17-4PH material exhibited greater strength and more elongation.

This can be attributed to the gaps present at the interface, formed due to surface roughness.Figure 11 Tensile stress-strain curve for brazing pairs.Table 5 Tensile stress-strain values of brazing pairs.Brazing Pair Tensile Stress (Mpa) Tensile Strain (%)C304- AM1742600,2C174-AM1742250,1C174-C1742690,1C304-C1742890,5C304-C3042901AM174-AM1742100,1ConclusionThe aim of this study was to investigate the vacuum brazing process using additive manufactured 17-4 PH material, conventional 17-4 PH material, and conventional AISI304 material. For this purpose, the vacuum brazing process was performed with these materials and material combinations under the same vacuum brazing thermal cycle. In this study, differences were observed through wetting analysis, microstructure analysis, and mechanical testing, and the study was completed.

In this study, differences were observed through wetting analysis, microstructure analysis, and mechanical testing, and the study was completed. Based on the obtained data, the following conclusions can be made:359Wetting analysis showed that the wetting index of parts manufactured by additive manufacturing was significantly higher due to the surface roughness. As the surface roughness increased, the wetting tendency of the brazing filler material also increased up to a certain point.Although the wetting of the brazing filler material on the surfaces of materials produced by additive manufacturing was high, it was observed that the diffusion within the material was lower compared to AISI 304.Microstructure analysis revealed that the BNi-2 filler material diffused more into the conventional AISI 304 material. The high carbon content in the composition of the AISI 304 material resulted in the formation of carbides both within the base material and in the brazing region.

The high carbon content in the composition of the AISI 304 material resulted in the formation of carbides both within the base material and in the brazing region. When brazing pair with 17- 4 PH material and AISI 304 material, these precipitates were observed to be closer to the AISI304 side. In experiments using the AISI304 material, the precipitates exhibited a larger and heterogeneous distribution.Surface roughness led to the observation of voids at the brazing interface between the additive manufactured 17-4 PH material and the brazing filler. These voids were found to reduce tensile stress and strain by acting as stress concentration points.The conventional 17-4PH material exhibited similar voids in the microstructure of brazing joints due to the oxide layer remaining from the solution treatment on its surface. These voids also had a negative effect on the mechanical strength.References[1] Davis, J. R. (Ed.). (2005). Stainless Steels for Design Engineers.

Used to reduce thermal distortion and brace large overhands, support structures often require post-print machining operations to remove, adding costs and processing delays. This preliminary work demonstrates that soluble, sacrificial metal supports can be fabricated in DED and PBF printers by taking advantage of small differences in the chemical and electrochemical stability between different metallic alloys. For DED printing, we demonstrate this process by printing stainless steel bridge structure with 90˚ overhang and printed carbon steel acting as a sacrificial support. For PBF printing, a PBF printed stainless steel part was first printed and then carburized to reduce the free chromium at the surface. Since the support/component interface is only ~100 μm in size, this interface becomes highly susceptible to chemical and electrochemical dissolution. In both cases, the component was separated from the supports in a solution of nitric acid and KCl under mild electrochemical bias.

In both cases, the component was separated from the supports in a solution of nitric acid and KCl under mild electrochemical bias. No machining, grinding, or finishing operations were required to remove the metallic supports. These novel approaches introduce new capabilities to additive manufacturing that will drastically reduce the post-processing needed for 3D printed metal components.Introduction Additive Manufacturing (AM) of metal components can be classified by their binding method (sintering, melting, polymer adhesive), energy delivery method (laser, electron beam), and metal feed method (powder- bed, powder-fed, wire-fed). Metal components are commonly built using one of two methods: Powder-Bed Fusion (PBF) or Directed-Energy Deposition (DED).1 In PBF, a laser or electron beam is scanned over a bed of metal powder to locally sinter or melt the powder, forming a slice of a part.

Building up a sequence of slices by with more powder and sintering layers forms a complete 3D metal component.1 In DED printing, metal powder(s) are blown or a wire is fed into a melt pool formed by a laser or electron beam. This added material increases the melt-pool and parts are built up in a layer-by-layer manner by moving the substrate relative to the energy/material deposition head.Components printed using PBF and DED printers often require support structures to minimize thermal distortion and form overhanging structures.2,3 Unfortunately, removing these support structures can require extensive post-process machining; which can add enough cost and processing time to outweigh any benefits associated with additive manufacturing. The plastics extrusion 3D printing community has already solved this problem by printing parts and supports out of polymers with different solubilities.

The plastics extrusion 3D printing community has already solved this problem by printing parts and supports out of polymers with different solubilities. For example, components are printed using Acrylonitrile Butandiene Styrene (ABS) and sacrificial supports printed using Polylaic Acid 1604(PLA). After printing, the PLA is selectively dissolved by immersing the printed structure in a solution of isopropyl alcohol and potassium hydroxide with the ABS component left intact.4This manuscript summarizes new advances in dissolvable metal supports for AM manufactured metal components. Using DED printing, we recently demonstrated that the concept of selectively soluble support dissolution can be applied to 3D printed metal structures.5 Specifically, we demonstrate that AISI type 431 (Fe 16Cr 2Ni 0.2C) stainless steel parts with free standing bridge structures with overhangs of 90˚ can be fabricated using Metco 91 (mild steel) carbon steel as the sacrificial support structure.

This support structure was removed using a solution of nitric acid with an applied external electrochemical bias to increase removal rate. While this set of experiments validated the concept of soluble metal supports, only a small fraction of the 3D metals printing community has access to multi-material DED printers. Instead, the community largely uses PBF printers that only print one material at a time. In this manuscript, also demonstrate a new process that brings dissolvable supports to single-material printed components. In this new process, we degrade the chemical resistance of the top ~100 μm of the printed structure using simple surface treatments incorporated into the thermal stress relieving step. Since support support/component interface is only 100-200 μm in diameter, the component is readily separated from the supports using chemical or electrochemical methods.This work draws heavily from the concept of sacrificial anodes.

A sacrificial anode is a material with a more negative redox potential relative to the part material such that the sacrificial anode will be preferentially Zn = -0.76 V) 7 is oxidized over the part material.6 For example, a zinc anode (standard reduction potential, Eo often used to protect stainless steel boat hulls since the standard reduction potential of Zn is lower than both the (cid:1)(cid:2)(cid:1)(cid:1)(cid:1) = 1.3 V)7 that forms in stainless base iron (Eo steel. As long as the zinc is in electrical and electrolytic contact with the stainless steel hull, then the zinc will preferentially oxidized to (cid:1)(cid:2)(cid:1)(cid:2) on redox potential of the constituent elements, but also microstructure, passivation, and electrolyte composition.8-15 As a result, small changes in composition or microstructure can result in very large differences in corrosion resistance.

This manuscript presents two different methods to generate composition and microstructure differences between the support and component part such that the support structures are preferentially etched. The first method used a multi-material DED printer to print dissolvable carbon steel supports for a stainless steel component. The second method used a single-material PBF printer to print a stainless steel part that is next sensitized using carburization/nitridation to captures the chromium in carbides/nitrides and lowers the corrosion resistance of the stainless steel to a depth of ~100 μm. Since the support structures only contact the part across 100 μm, the supports are easily dissolved away from the component.Fe = -0.44 V)7 and the passivating chromium oxide layer ((cid:1)(cid:1)(cid:1)(cid:1) instead of the steel.

For metallic alloys, the corrosion potential depends not justFor DED, the proof-of-concept system comprised of stainless steel as a component material, carbon steel asa sacrificial anode support material, and 41 wt.% nitric acid as the corrosive electrolyte. Stainless steel has excellent resistance to nitric acid while carbon steel is rapidly chemically dissolved nitric acid with or without an external bias.8,16,17 A free standing stainless steel bridge was fabricated using a DED powder-fed process by printing a stainless steel bridge with a carbon steel support across the middle and then etching away the carbon steel in nitric acid with bias of 700 mV to 900 mV relative to a the Saturated Hydrogen Electrode.Since multi-material PBF printers are not yet widely available, a surface-treatment approach for dissolvablesupports was developed for use in single-material systems. For a demonstration, a component part with attached supports was printed from PH 17-4 stainless steel.

For a demonstration, a component part with attached supports was printed from PH 17-4 stainless steel. After printing, the part was packed with sodium ferrocyanide hydrate and was heated to 800 ˚C for six hours. This treatment both reduced internal thermal stresses and diffused carbon and nitrogen into the top 100 μm to 200 μm of the PH 17-4. At this temperature, the carbon and nitrogen capture much of the available chromium in chromium carbides and nitrides. Once the free chromium content drops below 10.5 wt.%, the material is no longer “stainless” and readily dissolves under bias in nitric acid solutions. The supports were separated from the component at -150 mVSHE bias in a solution of 3 wt.% nitric acid and 0.1 molar KCl.This work is the first to demonstrate dissolvable support strategies for both multi-material printing systems(e.g., DED) and single-material printing systems (e.g., PBF).

These sacrificial materials will eliminate or reduce the need for post-process machining operations in 3D printed metals. Since both processes are “self-1605terminating,” they are extremely easy to implement and should free the design of AM fabricated metal components from the constraints of providing access to machining tools.Experimental MethodsDED Sample FabricationDED samples were fabricated using an Optomec Laser Engineered Net Shaping (LENs) MR-7, directed energy deposition system with two powder feeders was used for metal deposition. The system was equipped with a 500-watt Ytterbium-doped fiber laser (IPG YLR-500-SM). A second-moment beam spot size of 1.2 mm was measured, using a PRIMES GmbH FocusMonitor, at the substrate location. Before deposition, the substrate was positioned at a working distance of 9.3 mm from four, radially symmetrically powder-delivery nozzles.

Before deposition, the substrate was positioned at a working distance of 9.3 mm from four, radially symmetrically powder-delivery nozzles. AISI type 431 (Fe 16Cr 2Ni 0.2C) and Metco 91 (mild steel ) structures were deposited atop 304 Machinable Stainless Steel substrates. Prior to support dissolution, substrates were mounted in nitric acid resistance epoxy (Epoxy Systems, Inc., 633 Grey), cleaned with acetone and isopropyl alcohol and dried using N2 gas. See previous work for more detailed fabrication specifications.5All chemicals were used as received. The electrolyte solution was made by mixing 70 wt.% nitric acid (Sigma Aldrich, ACS reagent, 70 wt.%) with deionized water (18.2 MΩ, Purelab Flex 3) at a volume ratio of 1:1 – HNO3:H2O to form a final volume of 200 mL and a 41 wt.% nitric acid solution.

Next, 8 grams of KCl was added to improve the electrical conductivity of the electrolyte and help break down the passivation that forms on carbon steel parts in nitric acid.8PBF Sample Fabrication and SensitizationPH 17-4 samples were produced using an EOS M280 PBF system and built using the OEM’s standard parameters and powder for PH 17-4 (EOS StainlessSteel GP1) with a 20 μm layer height. Sample specimens were cut from a tubular arch with a measured OD of 8.16 mm, ID of 3.13 mm, and a major radius of 46 mm. The samples were sonicated in acetone, then isopropyl alcohol, and finally dried using N2 gas. Next, samples ⋅10H2O, Sigma Aldrich, were dipped in a saturated solution of sodium ferrocyanide decahydrate (Na4Fe(CN)6 ≥99%) allowed to air-dry, and placed on stainless steel tool wrap. A sodium ferrocyanide paste (5.4 grams ⋅10H2O, 1.3 mL DI H2O) was packed around the part and the tool wrap was folded and sealed Na4Fe(CN)6 around the part.

A sodium ferrocyanide paste (5.4 grams ⋅10H2O, 1.3 mL DI H2O) was packed around the part and the tool wrap was folded and sealed Na4Fe(CN)6 around the part. A small hole was punctured into the stainless steel wrap to outgas the ~0.25 moles of CN and H2O vapor that could be produced as the sodium ferrocyanide decahydrate decomposes. Samples were then place into an exhausted box furnace at 800 ˚C for six hours and allowed to cool to room temperature overnight in the oven. Samples were removed from the foil and excess decomposition products were removed using stainless steel bristle bushed followed by sonication in DI H2O and isopropyl alcohol.Electrochemical Analysis and DissolutionDED Support RemovalAll chemicals were used as received.

The electrolyte solution was made by mixing 70 wt.% nitric acid (Sigma Aldrich, ACS reagent, 70 wt.%) with deionized water (18.2 MΩ, Purelab Flex 3) at a volume ratio of 1:1 – HNO3:H2O to form a final volume of 200 mL and a 41 wt.% nitric acid solution. Next, 8 grams of KCl was added to improve the electrical conductivity of the electrolyte and help break down the passivation that forms on carbon steel parts in nitric acid. 8 Immediately prior to electrochemical analysis/dissolution, the samples were sonicated in acetone, then isopropyl alcohol, and finally dried using N2 gas.A Pine Research Instruments WaveNow USB Potentiostat/Galvanostat was used to measure the Open-Circuit Potential (OPC), gather potentiodynamic polarization curves, and apply constant bias for electrochemical dissolution of the mixed stainless steel/carbon steel part. All measurements were made relative to a saturated silver/silver chloride electrode in a saturated KCl solution.

All measurements were made relative to a saturated silver/silver chloride electrode in a saturated KCl solution. The reference electrode was interfaced to the etching electrolyte through a glass-fritted salt bridge with 4 M KCl (Alfa Aesar, 99%) intermediate electrolyte; the glass frit was located approximately 1.0 mm from the working electrode (i.e. 3D printed sample). The measured Potentials, E, were offset by +0.197 mV so that all reported values are relative to the1606Standard Hydrogen Electrode (SHE). A 6 mm diameter graphite rod was used as the counter electrode. The electrodes were positioned in a line with the reference electrode between the working and counter electrodes. PBF Support RemovalAn electrolyte of 3 wt.% nitric acid with 0.1 molar KCl was made by mixing 70 wt.% nitric acid with DI H2O, and KCl.

PBF Support RemovalAn electrolyte of 3 wt.% nitric acid with 0.1 molar KCl was made by mixing 70 wt.% nitric acid with DI H2O, and KCl. Using the same setup as the DED support removal process, OPC and potentiodynamic curves curves where gathered from the untreated part, treated part, and post-etched component to identify the corrosion potential and potential to selectively etch the sensitized surface without etching the component. The sample was etched between 99-444 mVSHE for 12.75 hours until the component separated from the part. The sample was removed every sixty minutes for optical imaging.

The sample was etched between 99-444 mVSHE for 12.75 hours until the component separated from the part. The sample was removed every sixty minutes for optical imaging. After eight hours, the build-up of black powder (presumably magnetite, Fe3O4) within the supports was removed using a steel wire brush and the sample was etched for two more hours until the component separated from the supports.Results and DiscussionDissolvable DED SupportsThe full details of the dissolvable DED support structures can be found in the author’s previous work.5 The corrosion potential, Ecorr, of the AISI type 431 stainless steel, Meco 91 carbon steel, and the mixed printed part were measured at 1.113 VSHE, 1.004 VSHE, 0.197 VSHE respectively. As expected, the stainless steel shows a more positive Ecorr than both the mixed sample and the carbon steel sample.

As expected, the stainless steel shows a more positive Ecorr than both the mixed sample and the carbon steel sample. More importantly, polarization curves showed that carbon steel etches many orders of magnitude faster than the carbon steel across wide range of potentials.5Electrochemical etching the mixed stainless/carbon steel sample was conducted by setting a constant potential (typically between ~800 mVSHE to 900 mVSHE) to create a slight anodic current, typically on the order of -10 mA to -50 mA, with no O2 bubbling but >100 mA with O2 bubbling. Since the Pinewave Now potentiostat is limited to ±105 mA and was often driven to maximum current, the actual applied varied with the formation of passivating layers, changing surface area, and bubble entrapment within the stainless steel bridge.

Under potentials between ~700 mVSHE to 900 mVSHE, a cyclic shift from anodic to cathodic with a periodicity between 50 to 500 seconds was observed as a passivation layer on the carbon steel forms and dissolves. 5 Under potentials between ~700 mVSHE to 900 mVSHE, a cyclic shift from anodic to cathodic with a periodicity between 50 to 500 seconds was observed as a passivation layer on the carbon steel forms and dissolves.5,18 Under these conditions, the stainless steel is well protected; however, the etch rate is too slow for practical applications, and it took ten hours to remove 1.4 mm of carbon steel from each end. The etch rate was increased significantly by bubbling O2 onto the carbon steel section of the part. The O2 helps break down passivated carbon steel, and a cathodic current was observed. This increases the etch rate dramatically, and the remaining 7 mm of the carbon steel section was removed in 6 hours.

This increases the etch rate dramatically, and the remaining 7 mm of the carbon steel section was removed in 6 hours. This constant cathodic current was not observed when bubbling O2 onto the stainless steel section of the sample. Figure 1 shows the sample before and after etching. Notice that the carbon support material was completely removed during the etching process. The support removal process is, effectively, self-terminating and should be robust for manufacturing applications.1607DED Dissolvable Metal Supportsa)Before Etchingb)After EtchingStainless SteelCarbon SteelFigure 1. Digital images of the sample taken before and after etching. (a) Before etch, red square roughly outlines where the carbon steel support layer was printed. (b) After etching showing that the carbon steel was completely etched from the sample.5Dissolvable PBF Support StructuresAs a demonstration, a simple tubular arch was printed from PH 17-4 stainless steel using PBF printer.

Sections of the part were cut from the tube (see Figure 2a) and subject to a sensitization processes to capture the free-chromium within the top 100 μm – 200 μm of the surface. While there are numerous ways to capture this chromium, this concept was demonstrated with a standard “packed” carburization process that used sodium ferrocyanide as the carbon and nitrogen source. The sample was packed with sodium ferrocyanide and treated for six hours at 800 ˚C. At this temperature, chromium carbide precipitates readily form at the grain boundaries19 and the available “free-carbide” drops below the 10.5 wt.% necessary to sufficiently passivate the part with protective chromium oxide where the carbon content was high enough.The sample was examined and etched in a 3 wt.% nitric acid solution with 0.1 molar KCl added to improvethe conductivity of the solution. The measured OPC of the untreated, sensitization treated, and post-etch component were 365 mVSHE, -22 mVSHE, 363 mVSHE respectively.

The measured OPC of the untreated, sensitization treated, and post-etch component were 365 mVSHE, -22 mVSHE, 363 mVSHE respectively. To dissolve off the supports, an anodic potential between 99 mVSHE to 444 mVSHE was applied for 8 hours with the sample removed every hour for optical imaging. During this time, the corrosion current was observed to decrease over time even with anodic potentials were applied. A buildup of black powder was observed over the sample and was suspected of restricting electrolyte access to the sample surface. A steel wire brush was gently applied to the sample surface followed by sonication, a dark black powder was observed coming off the sample. This process was repeated eight times until no more powder was observed. After this first cleaning process, the measured current returned to 80 mA. This verified that precipitate build-up was responsible for the drop in current, future work will add HCl to dissolve the precipitated manganite and avoid this build up.

This verified that precipitate build-up was responsible for the drop in current, future work will add HCl to dissolve the precipitated manganite and avoid this build up. For this work, the precipitate was removed using wire-brushing and sonication; although less extensive cleaning was needed for the remainder of the etching process. The sample was etched for another 3.5 hours before the component separated from the supports.The optical images in Figure 2a-f shows the sample as it when through this new process. Figure 2a shows the original printed sample. The “component” is the round tubular section on the left. Notice that the supports connect to the component at fine points less than 100 μm in size. After heat treatment, Figure 2b, the sample is darkened in color and some of the voids between the sample and the component have been filled in with excess carbon. Although not applied in this experiment, this excess carbon can be easily burned away to increase the initial etch rate.

Although not applied in this experiment, this excess carbon can be easily burned away to increase the initial etch rate. Figure 2c shows the sample after eight hours of etching with no brushing. Notice that the component still appears dark in color, indicating that the carburized layer has not been fully removed. Additionally, only a small percentage of the supports have been removed by this time. Brushing and sonicating the sample to remove precipitate buildup increases the etch rate significantly. Figure 2d-f show significant support dissolution over the 3 hours with component separating with no mechanical machining operations.1608Notice that by Figure 2d, the component appears bight and shiny while the support structure is still dark and black. This indicates that the etching has reached the underlying stainless steel. The diameter was measured after sensitization and hourly after eight hours.

This indicates that the etching has reached the underlying stainless steel. The diameter was measured after sensitization and hourly after eight hours. The original diameter was 8.16 mm after eight hours of etching the diameter decreased to 7.85 mm. By nine hours (after brushing) of etching the underlying stainless steel was revealed and the diameter was measured at 7.78 mm and did not decrease or change with further exposure to the etching environment; indicating that the process is self limiting and can be conducted without overly-detailed monitoring. Future work will eliminate the need for the brushing and sonication steps by adjusting the electrolyte to dissolve and reaction products.a) printed sampleb) after sensitizationc) etching - carbonurized layer still on componentd) etching - carburization layer removed from componente) etching - almost donef) separation completeFigure 2. Camera images of sample.

Camera images of sample. a) as printed; b) after sensitization treatment; c) after 8 hours of etching. Notice that the surface of the component (round section) is still black in color and that the etch rate appears slow when magnenite isn’t removed from the surface. d) after 9 hours (total) of etching with wire- brushing. Notice that the component is now shiny and bright; indicating that the stainless steel layer has been reached. e) after 10 hours (total) of etching. Notice that the component appears unchanged while most of the support structures have been dissolved; f) after 11 hours of etching. The component completely separated from the supports with no mechanical machining. Component appears largely unchanged.Conclusions In conclusion, a method to form large overhangs and dissolvable support structures in parts fabricated using DED and PBF additive manufacturing.

Component appears largely unchanged.Conclusions In conclusion, a method to form large overhangs and dissolvable support structures in parts fabricated using DED and PBF additive manufacturing. For DED, a stainless steel bridge was fabricated using carbon steel that was later removed by electrochemically etching the carbon steel in a 41 wt.% solution of nitric acid with bubbling O2. For PBF, a stainless steel sample was fabricated and then sensitized using a carburization process 1609that effectively converted the top ~100 μm of stainless steel to carbon steel. This layer was electrochemically etched in a 3 wt.% nitric acid solution with 0.1 molar KCl. The supports were easily separated with no mechanical machining operations. This first-of-its-kind approach introduces new capabilities to additive manufacturing and may drastically reduce the post-processing needed for metal parts thereby reducing process costs while simultaneously increasing design freedom.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDistributed Optical Fiber Testing for Additive ManufacturingBrian M. Hlifka, Edward C. KinzelDepartment of Aerospace and Mechanical Engineering, University of Notre Dame, Notre Dame, IN 46556AbstractThis paper explores optical fiber's use for in-situ inspection of additive manufacturing. Single-mode SMF-28 optical fiber can be placed on the build plate to monitor the printing process or embedded in the part. Distributed measurements using optical backscattering reflectometry (OBR) resolve the strain along the optical fiber and the temperature. OBR-enabled sensing is demonstrated for the fused filament fabrication (FFF) process. The small diameter (0.125 mm) of SMF-28 lends itself to embedding in FFF prints. This enables laying the fiber into the part, which provides continued sensing for the details in use.

The small diameter (0.125 mm) of SMF-28 lends itself to embedding in FFF prints. This enables laying the fiber into the part, which provides continued sensing for the details in use. Knowledge of the process and the fiber arrangement allows heating from the deposition head to be distinguished from stress-driven strain. Calibration of the fiber arrangement is discussed, as well as a comparison with process modeling.1. IntroductionOptical fibers have widespread applications in long-distance telecommunications, optical imaging, and advanced optical sensors [1]. This versatility is due to the material’s designed optical properties (low losses over long distances), immunity to electromagnetic interference, lightweight construction, compact dimensions, and the ability to accommodate distributed sensor systems [2].

Consequently, distributed fiber sensors are increasingly employed for real-time monitoring of in- process and fabricated materials, providing distinctive capabilities for quality control and performance assessment along the length of the fiber [3]. These distributed sensors can map various physical parameters, including temperature, strain, acoustic vibrations, electric current, and pressure, by detecting changes in light propagation and scattering that result from localized strain profiles along the fiber [4]. This capability to capture strain (both mechanical and by thermal expansion) distributed along the length of the fiber enables in-situ monitoring of a process, but the physical design of the fiber offers additional benefits, particularly in the field of additive manufacturing (AM).

Making subtle changes to the design of AM parts allows for optical fiber to be mechanically embedded into the part, which can be used for in-situ recording of the deposition temperature / strain fields. Additionally, the fiber remains mechanically bonded to the part after production allowing for real-time health monitoring of the part in the field, and the extremely small diameters of fibers (0.125 mm) limits the impact on the part’s mechanical properties.The integration of distributed optical fiber sensors into manufactured components, particularly those composed of composite materials, has yielded success. Early endeavors in this domain primarily revolved around Fiber Bragg Gratings (FBGs), which exhibit a longitudinally periodic variation in refractive index within the core of single-mode optical fibers [4].

When employed as embedded sensors within composite materials, FBGs have proven valuable for in-situ measurements across a spectrum of parameters, encompassing strain, temperature, corrosion, vibration, and even the state of cure [5]. However, it is worth noting that the fabrication of FBGs can often be cost-prohibitive due to the intricate microstructure creation within the fiber core, and1055their utility is limited to discrete point measurements. Consequently, recent developments have shifted the focus toward deploying distributed optical fiber sensors (DOFS) in fabricated materials, enabling measurements distributed across a manufactured part. DOFS utilizes the innate structural irregularities within optical fibers to capture Rayleigh scattering phenomena along the fibers’ length, which can be calibrated to detail temperature and strain measurements.

The origin of Rayleigh scattering is related to the refractive index fluctuations caused by density and compositional inhomogeneities in the fiber structure created during the fabrication process when the temperature cools from its glass softening point back down to room temperature [6]. As illustrated in Figure 1, light can travel great distances in two primary directions due to the total internal reflection at the core-cladding interface, which is forward and backward. The Rayleigh backscattered light forms the basis of operation for optical backscattering reflectometry (OBR) measurement techniques.Fiber Claddinge r o CInput LightIncident Lightr e b FiBackscattered LightFiber CladdingFig 1. Schematic illustrating Rayleigh scattering phenomenon in optical fibersFurthermore, the distinctive attributes of DOFS make them highly pertinent to the evolving AM landscape as a quality control measurement technique [7].

The integration of DOFS has garnered the context of Fused-Filament-Fabrication (FFF) substantial attention, particularly manufactured components. This integration provides AM structures with the capability of real- time monitoring of temperature and strain variations, thus providing crucial insights into the performance and structural integrity of FFF-fabricated objects [8,9]. Recent investigations have delved deeper into optimizing coating adhesion and channel depth techniques for embedding optical fiber sensors within FFF components [10]. These advancements elevate the role of DOFS in AM, pave the way for innovative applications, and enhance quality control standards for applications beyond FFF.inThe primary aim of this research article is to formulate a comprehensive set of guidelines geared towards accurately measuring temperature and strain within FFF components.

These guidelines are intended to serve as a foundational reference and facilitate integration into complex geometries. The approach will offer insight into the constraints and capabilities associated with integrating DOFS in both AM components and in-process workflow, thereby better understanding their practical utility and limitations. Notably, a Warren truss structure will be employed as an illustrative example. The knowledge derived from these methodologies will be actively used for continuously monitoring strain fluctuations throughout the entirety of the FFF manufacturing process.2. Measurement Method The experimental DOFS measurements were performed using Optical Backscattering Reflectometry (OBR). This method commonly employs a tunable laser source (TSL) that emits a continuously changing wavelength within the c-band spectrum over time. The light emitted by this laser source is split into two paths: one measurement and the other for reference.

The light emitted by this laser source is split into two paths: one measurement and the other for reference. In the measurement path, the light is coupled into the optical fiber sensor, where it interacts with the1056optical fiber’s natural inhomogeneities, leading to Rayleigh backscattering. This backscattered light serves as the returning signal. Conversely, in the reference path, the light is directed to a beam splitter, where the scattering profiles from both the measurement and reference arms are collected. Subsequently, these two scattering profiles are cross-correlated to assess their similarities, primarily based on the shifts observed. This correlation process translates into measurements of strain and temperature using calibration coefficients that have been empirically determined [11,12,13].An ODiSI 6100 series interrogator was utilized for all distributed optical fiber measurements based on OBR.

This interrogator measures the recorded spectral shift responses, specifically from the intrinsic Rayleigh backscattering signal within an unaltered standard telecommunications-grade optical fiber. These recorded shifts corresponded to variations in both temperature and strain. The concept of spectral shift is described as follows:∆𝜆𝜆=∆𝑓𝑓= 𝐾𝑇∆𝑇 + 𝐾𝜀∆𝜀In this equation, KT represents the sum of the thermal expansion and thermo-optic coefficients. Typical values for these coefficients in Germania-doped silica core fibers are 0.55x10-6°C-1 and 6.1x10-6°C-1, respectively. The strain coefficient, Kε, depends on factors such as group index, strain optic sensor components, and Poisson ratio, resulting in a value of 0.787 [14]. Consequently, when measuring temperature, the strain coefficient is disregarded, allowing for the measurement of spectral shifts using the known constant to determine temperature and vice versa with strain.3.

Optical Fiber Sensor Integration in AMThis section provides the guidelines for effectively integrating DOFS into PLA printed specimens through the FFF process. Essential integration techniques for temperature and strain- based measurement applications will be analyzed. All embedded test specimens were produced using the Prusa i3MKS 3D printer, employing a 1.75mm PLA feedstock for each print. Standard printing parameters were maintained, including a nozzle temperature of 215°C and a heated build plate set at 60°C. The layer raster angle was fixed at 45°, and the layer height was set at 0.20mm.3.1.1. Embedding Optical Fiber in the Warren TrussIn the case of each embedded PLA specimen, printing was paused at a specified height. This pause allowed for the precise placement and integration of the optical fiber into the print. This precision is crucial, as the quality of the integration and resulting signal depends on proper placement security to the part.

This precision is crucial, as the quality of the integration and resulting signal depends on proper placement security to the part. Therefore, the optical fiber was laid out on the top surface of the sectioned part and was initially secured using Kapton tape. The slack was then removed from this initial layout. External printed jigs were also used to simplify the processes aligning the optical fibers in a known orientation. Once the optical fiber is secured with tape, cyanoacrylate glue is applied to bond the fiber and part together. The epoxy was left to set for ~10 minutes before the printing process continued. The effect of cure duration is not characterized in this study.To illustrate the capabilities of the integrated optical fiber sensors, a modified Warren truss structure was selected as the representative sample for experimentation.

The Warren truss offers more geometric complexity than a tensile test (dog bone) specimen, providing an excellent platform for monitoring temperature and strain variations within a nontrivial design. The standard Warren truss design was slightly modified to allow for the DOFS implementation, incorporating1057(1)of a 20mm bending radius to prevent any light leakage from the fiber (refer to Figure 2). When optical fibers are bent excessively, light may escape through the fiber's cladding, which hinders the interrogator's ability to obtain accurate measurements [15].(2)2 mm4 mm2 mm20 mm60 mm(3)L02 mm2 mmL1(1)Fig 2. Design of modified Warren truss structure with internal beam bending radius of 20mmAs shown in Figure 2, the SMF-28 fiber was laid out in three geometries over the cross-section of the truss shape.

Design of modified Warren truss structure with internal beam bending radius of 20mmAs shown in Figure 2, the SMF-28 fiber was laid out in three geometries over the cross-section of the truss shape. The layout was specifically selected to allow for simultaneous measurements of all members of the shape, while still maintaining full density in the sections of the design with the largest stress concentrations. Each member of the truss was designed to be a 4  4 squares with the fiber placed in the center (2 mm above the heat bed, 2 mm from the edge of the part). One of the initial challenges that appeared with using this method was the registration of the fiber embedded in the part compared to the rest of the sensing fiber. To address this registration issue, after the part had been printed, the ODiSI system was used to measure the temperature in the fiber while the tip of a soldering iron (~50°C) was placed at the ends of the embedded fiber (L0 and L1 in Figure 1).

The ODiSI system detected these temperature changes, with the length of fiber in- between them defining the lengths of embedded fiber.It is worth noting that since the DOFS technique is designed to measure changes in strain, there can be significant challenges with deconvoluting a combined mechanical – thermal strain profile without a thorough understanding of the interaction between these profiles. This is further complicated by external interactions such as friction between the sample and fiber, as well as the thermal conductivity between the materials. As such, we devised two experimental tests to highlight the design decisions to optimize a produced part for a particular type of strain, aiming to identify the trade-offs between the two methods.For measuring the mechanical strain internally, it is imperative that the fiber be mechanically well bonded to the surrounding material such that the forces can be effectively transferred to the fiber.

As such, this design did not incorporate any designed channels throughout the bulk of the part, instead relying on the adhesive and contact to the part to hold the parts in place. However, to prevent excessive shear stress on the fiber at the edges of the part, a 111 mm channel was designed at the edge of the part for strain relief. Note that since there are no internal channels designed through the bulk of the part, the fiber must be aligned effectively by hand for embedding into the part, however this can be improved over time with better jigging.1058When measuring the temperature profile inside the part, some additional considerations must be made in the design.

When the fiber is very strongly bonded to the surrounding material, a change in temperature will not only modify the length of the fiber (thermal strain), but the coefficient of thermal expansion (CTE) mismatch between the fiber (0.5  10-6 ε/°C) and PLA (68  10-6 ε/°C) is sufficiently large to introduce significant thermal stress in the fiber. This additional tensile stress may compound with the thermal strain, leading to significant errors in the temperature measurements [16,17]. To avoid this phenomenon, the design must be altered such that the fiber is not strongly bonded to the part. This is done by incorporating 11 mm channel into the truss (following the desired path displayed in Figure 1). These pre-built channels allow for significantly more accurate placement of the fiber inside the part.3.1.3.

These pre-built channels allow for significantly more accurate placement of the fiber inside the part.3.1.3. Distributed Strain MeasurementsTo demonstrate the ability of the DOFS method to capture the internal mechanical strain profile, the samples were tested in a Mark-10 Universal Testing Machine (UTM), in a 3-point bending configuration (Refer to Figure 3).Fig 3. Embedded warren truss structure during three-point bend test with incrementally increasing loadLoads of 10N, 20N, and 30N were applied, showing a linear displacement increase on the top chord of the Warren truss structure based on the in-situ measurements from the optical fiber sensor (see Figure 4). Additionally, the in-situ measurements were compared to simulations conducted using ANSY. It’s important to note, that the optical fibers are sensitive to normal strains, therefore, direct comparisons from the top and bottom chord were made based on their given simplified orientations.

It’s important to note, that the optical fibers are sensitive to normal strains, therefore, direct comparisons from the top and bottom chord were made based on their given simplified orientations. Also, 3D-viz software was used to visualize the in-situ strain data from the optical fiber sensor distributed along the Warren truss structure. The findings demonstrate a strong correlation between the measured data and the results obtained through ANSYS simulations; however, there are discrepancies for the measured tension in the beam members (see Figure 5).105910N30NStrain [me]Position [mm]Top Chord20N3045607590−800−600−400−2000Fig 4. Linear strain response to increasing load on the top chord of the Warren truss25µεA)150µεB)Fig 5.

Linear strain response to increasing load on the top chord of the Warren truss25µεA)150µεB)Fig 5. A) In-situ strain response to a 10N applied force in a three-point bend apparatus compared to ANSYS resolved strains for the B) top chord and C) bottom chord of the Warren truss structureThe ANSYS-simulated data, when subjected to a 10N applied load, displays a notable correlation concerning compressive loads. However, there is a substantial inaccuracy in the reflected tension measurements. This discrepancy can primarily be attributed to the placement of the optical fiber within the beam members. Achieving precise optical fiber placement is challenging, particularly when optimizing for strain-related applications, as there are no designated channels for securing the optical fiber sensors within the beam members.C)Precision in placement plays a pivotal role in obtaining measurements that closely match simulation results.

In this context, each optical fiber was intended to be positioned at the neutral axis of each member, where the beam’s width measures 4mm. However, achieving this ideal placement posed challenges and uncertainties, particularly at the edges of the structure. The primary concern with optical fibers structure’s edges is the potential for displacement or increased stress due to overlap, which can result in issues such as light leakage or fractures. There were instances in the Warren truss structure where optical fibers overlapped with the internal structural1060beam members and the bottom chord, presenting challenges as the optical fibers became susceptible to breakage or displacement during the printing process.To address this issues, 1mm x 1mm x 1mm channels were introduced at the edges of each structural member.

These channels serve a dual purpose: they permit optical fibers overlap and ensure precise placement within the structure without compromising their sensitivity to strain. It’s worth noting that the use of these channels at the edges introduce minor variations in strain measurements between the edges of each optical fiber, potentially affecting the resulting strain measurements within the truss structure. Moreover, concerns may also arise regarding the effectiveness of the cyanoacrylate glue in bonding the optical fiber to the PLA specimen. A need exists for more comprehensive and in-depth investigations to gain thorough understanding of how these channels, and choice of adhesive can influence the resulting strain measurements.3.1.3. Distributed Temperature MeasurementsIn addition to investigating strain measurements, this study also delved into optimizing the warren truss structure for temperature monitoring.

Distributed Temperature MeasurementsIn addition to investigating strain measurements, this study also delved into optimizing the warren truss structure for temperature monitoring. For this purpose, a thermal camera was employed to capture temperature variations across the Warren truss structure. This enabled comparison from the thermal data obtained through the camera with the distributed measurements obtained from the optical fiber sensor.To optimize the Warren truss structure for temperature measurements, one critical aspect was considered, which was ensuring the optical fiber's precise placement within the structure. This optimization process included channel depths of 1mm x 1mm along each structural member.

This optimization process included channel depths of 1mm x 1mm along each structural member. These channels played a crucial role in providing the optical fiber with sufficient space, mitigating the potential challenges arising from the differences in coefficients of thermal expansion (CTE) between the silica cladding of the optical fiber and the PLA material used in the truss structure.To facilitate temperature measurements within the Warren truss structure, an Optris PI 640i infrared camera was employed. This camera allowed precise temperature data to be captured across the truss structure, providing a valuable benchmark for comparing the distributed temperature measurements obtained from the embedded optical fiber sensors. Also, adjustments to the emissivity were changed to 0.92 to match the molten polymer [18].Controlled temperature experimentation was conducted to facilitate a comparison between the thermal camera and the DOFS.

The embedded Warren truss structure was exposed to an external temperature source generated by a Corning PC-400D digital hot plate. A predetermined temperature of 50°C was established as the reference point for this comparative analysis. The placement of the Warren truss structure on the hot plate was strategically chosen to accentuate the temperature variations within the structure. By positioning and securing the truss structure at the center of the hot plate, it was ensured that the central region would experience the highest temperature, with temperatures diminishing towards the outer portions of the structure. This arrangement aimed to provide a clear demonstration of how temperature gradient would manifest across the truss.1061Like the approach taken for strain measurements, temperature changes were monitored using a known temperature source, such as the thermal camera, while the distributed optical fiber measurements were visualized and analyzed using 3D-viz software.

The results, presented in Figure 6, reveal a significant level of correlation between the temperature data obtained from the thermal camera and the optical fiber sensor.50.3°C22°C50°C22°CTCDOFS05010015020020253035404550T[°C]Position [mm]Fig 6. Temperature response comparison between the thermal camera and embedded optical fiber sensors whenexposed to 50°CThe results obtained from the temperature optimization highlight the enhanced capabilities of the integrated optical fiber sensor for temperature measurements within complex structures like the Warren truss. Incorporating embedment channels significantly improves measurement accuracy compared to a Warren truss structure optimized for strain sensing. It is essential to note that the differences in the CTE between the optical fiber and PLA introduce additional strain changes within the part as the temperature fluctuates.

It is essential to note that the differences in the CTE between the optical fiber and PLA introduce additional strain changes within the part as the temperature fluctuates. The fundamental principles underlying the ODiSI 6100 series interrogator remain consistent, as all temperatures and strain measurements rely on spectral shifts in Rayleigh scattering sites transitioning from their initial state to a stressed state1062induced by temperature or strain variations. These variations in measurement accuracy are exemplified in the results obtained from the two fabricated truss samples exposed to a temperature of 50°C (Refer to Figure 7).02040608010001020304050T[°C]Position [mm]Top ChordCompactLooseBottom Chord050100150200050100150200T[°C]Position [mm]050100150200250300010203040506070T [°C]Position [mm]Internal BeamsFig 7. Temperature measurements in embedded Warren truss samples illustrate two interface conditions: compact and loose when subjected to a temperature of 50°C.

Temperature measurements in embedded Warren truss samples illustrate two interface conditions: compact and loose when subjected to a temperature of 50°C. The loose interface sample features structural channelsThe Warren truss sample that was fabricated without the use of channel exhibits superior bonding to the specimen when compared to the fiber placed with the designated channel. Consequently, the difference in CTE for the PLA affects the resulting temperature measurement in the strain optimized sample. As the temperature rises, the PLA material expands, leading to a shift in Rayleigh backscattering that simultaneously measures both temperature and strain. Therefore, there are ongoing challenges in applying optical fiber to accurately capture both temperature and train measurements. This is because the ODiSI system measures a pure shift of Rayleigh scatter and correlates it to ither a strain or temperature, the system cannot simultaneously measure both.

This is because the ODiSI system measures a pure shift of Rayleigh scatter and correlates it to ither a strain or temperature, the system cannot simultaneously measure both. As a result, there are inherent discrepancies in measurements based on the specified application.4. Integration with the FFF ProcessThe refined approach for strain measurement has been successfully applied to more intricate structures, such as the warren truss. The insights gained from this experience can be directly implemented within an additive manufacturing (AM) process. To illustrate this capability, the placement of an optical fiber on the build plate has been demonstrated, showing its potential to monitor strain and temperature fluctuations. Masking tape secured the optical fiber, ensuring its stability and proper positioning throughout the FFF printing process (refer to Figure 8).1063Fabricated PartMasking Tape SMF-28Build PlateFig 8.

Prusa i3MKS printer and schematic depicting the configuration of the optical fiber on the build plateWhen subjected to an applied thermal load resulting from the heating and cooling of molten PLA material, the optical fiber is sensitive to thermal and strain changes from the part. A 200mm x 12.7mm test specimen was printed with the optical fiber aligned along its neutral axis to establish a baseline measurement for incorporating the optical fiber onto the build plate. The spectral shift was measured for each layer, observing a decreasing negative shift peak after each subsequent layer (refer to Figure 9). The test print adhered to standard printing conditions, including a raster angle of 45° and a layer height of 0.20mm.10641 2 . 7 m mx200 mm050100150200−120−100−80−60−40−20020406080Spectral Shift [Df]Position [mm]Layer 1Layer 4Layer 3Layer 5Layer 2Layer 6Fig 9.

7 m mx200 mm050100150200−120−100−80−60−40−20020406080Spectral Shift [Df]Position [mm]Layer 1Layer 4Layer 3Layer 5Layer 2Layer 6Fig 9. Spectral shift response for layer-by-layer deposition for a 200mmx12.7mm test specimenThe optical fiber response stabilizes after approximately five layers, equivalent to a height of 1mm. However, there is a limitation to integrating the optical fiber on the build plate, as its sensitivity diminishes progressively with the increasing distance from the build plate due to the reduced thermal load. Nevertheless, the optical fiber continues to effectively detect external shear forces resulting from the interaction between the nozzle and the printed part. The optical fiber’s sensitivity to normal forces along its designated position on the bed allows it to perceive shear forces induced by the nozzle interaction with the part.

The optical fiber’s sensitivity to normal forces along its designated position on the bed allows it to perceive shear forces induced by the nozzle interaction with the part. This aspect is intriguing because regardless of the printed part’s height, the induced shear force remains constant, as it is a relationship between the force and area of the print.𝜏 =𝐹𝐴To verify the consistency of shear forces, a test specimen measuring 12.7mm x 12.7mm x 38.1mm was printed, where the optical fiber was secured to the build plate along the neutral axis of the part. This test aimed to demonstrate that a constant shear force of 5N remained uniform, irrespective of the height at which it was applied (refer to Figure 10).

This test aimed to demonstrate that a constant shear force of 5N remained uniform, irrespective of the height at which it was applied (refer to Figure 10). The forces were applied using a fish scale, and a soldering iron was employed to create an area for the fish scale to exert force, resulting in an approximate 5N force.1065(2)FFPrinted PartHx𝜀 xPrinted PartM𝜀 yMasking TapeSMF-28Build Platey12.7 mm38.1 mm051015202501020304050Strain [me]Y Position [mm]25.4 mmFFig 10. Illustration of shear forces applied to a printed component and corresponding strain measurements for an applied force of approximately 5NThe resultant shear forces from an applied load are the same despite the height where the force is applied. However, the amount of force required to remove the printed part decreases as the height increases for the part.

However, the amount of force required to remove the printed part decreases as the height increases for the part. It's important to note that before the removal of the part with the inclusion of masking tape, the printed piece is attached to the tape, so with an increasing shear force, there will be an equal and opposing force in parallel to the direction of the load and an opposite normal force, where a spike in strain will be shown before completely being ripped off the build plate. This phenomenon was exemplified using the same printed specimen. However, a load of 12.8N was applied at a height of 12.7mm, and a load of 14.2N was applied at 25.4mm. The results reveal a strain profile with significant peaks corresponding to the opposing forces acting on the printed part (see Figure 11).1066M12.8N14.2 N0255075100125150−100−50050100150200250Strain [me]Position [mm]Fig 11.

Strain response of optical fiber to external loads at different heights with varying forcesEven with a lower load at greater heights, a substantial opposing force is generated to retain the part of the build plate. This extends the utility of integrated optical fiber beyond the initial few printed layers. Nonetheless, further comprehensive investigations are required to enhance its robustness. These studies could establish a correlation between the measured strain response and the point at which potential detachment from the bed or failure might occur.5. ConclusionsOptical fiber sensors have demonstrated their versatility as invaluable tools for real-time monitoring in additive manufacturing (AM) processes, providing critical data on temperature and strain. Their unique capacity to capture distributed physical information along the length of the fiber offers valuable insights into material behavior while maintaining the structural integrity of the components.

Specifically, distributed optical fiber sensors (DOFS) have emerged as promising alternatives to traditional discrete point sensors like Fiber Bragg Gratings (FBGs). DOFS enables measurements to be distributed across the entirety of the manufactured part, thereby enhancing the capabilities for quality control and performance assessment in AM components.Therefore, ensuring the precise placement and secure attachment of optical fibers within AM components is imperative to guarantee accurate measurements. The optimization of temperature measurements within the Warren truss structure exemplifies the enhanced potential of integrated optical fiber sensors. Incorporating embedment channels has notably elevated measurement accuracy compared to structures primarily tailored for strain sensing.1067Moreover, further research endeavors are warranted to strengthen the robustness of optical fiber integration within AM processes.

DITHERING BASED METHOD TO GENERATE VARIABLE VOLUME LATTICE CELLS FOR ADDITIVE MANUFACTURINGD. Brackett, I. Ashcroft, R. HagueWolfson School of Mechanical and Manufacturing Engineering, Loughborough University, Loughborough, Leicestershire, LE11 3TU, UKAbstractThis paper covers the principles of a novel method to efficiently spatially vary the size of tetrahedral cells of a lattice structure, based upon finite element analysis stress results. A dithering method, specifically error diffusion, is used to represent a grayscale stress fringe with variably spaced black dots. This enables linkage of the spacing between lattice cell vertices to stress level thereby providing a functional variation in cell density. This method is demonstrated with a simple test case in 2D and the steps involved for extension to 3D are described.IntroductionLattice structures generally contain a large number of cells, each made up of several structural members.

They can be used to control the stiffness of a structure or to provide tailored impact absorption capacity, usually through plastic deformation. Additive manufacturing (AM) is more suited to the manufacture of these complex structures than traditional manufacturing processes. The design of these structures, however, remains a challenge. With increased geometric complexity comes increased design complexity, and this is exacerbated when including computational analysis methods and mathematical optimization techniques in the design process. Some approaches focus on how to handle the geometric complexity problem while others focus on how to handle the analysis/optimization complexities. Both of these approaches should be unified to make a useful lattice design tool.Different regions of a component generally require lattice support to different extents. With a fixed lattice, this variation can only be achieved by varying the structural member dimensions.

With a fixed lattice, this variation can only be achieved by varying the structural member dimensions. This leads to a large number of design variables, depending on the resolution required. In these cases, the variation in support cannot be tuned by adjusting the sizes of the lattice cells and so the end result would be expected to be less than optimal due to the design freedom restrictions. Some approaches combine different cell structure designs together, which is difficult, and the cells are kept the same size [1]. Optimizing the size of each cell in a lattice is possible, but is computationally very expensive and awkward to implement, and so this approach is generally avoided. Some approaches have used different sized cells, not for structural performance reasons, but only as a by-product of requiring the lattice structure to conform to a shape with curved faces. Instead of trimming tessellated cells to this shape, they are either swept [2] or based upon an unstructured mesh [3].

Instead of trimming tessellated cells to this shape, they are either swept [2] or based upon an unstructured mesh [3]. Spatially varying the sizes of the lattice cells based upon structural analysis in an efficient manner is the topic of this paper.The method in this paper uses a dithering or halftoning method to define spatially varying points which are subsequently connected with lattice members. The points determine the spatialvariation of the lattice cell volumes. This approach was inspired by previous work on linking dithering to meshing techniques by [6-9].MethodThis section begins with a brief introduction to dithering before moving onto specific method details. Dithering is a procedure that converts continuous tone images into a binary representation. This is useful for bi-level printers and displays. When viewed from a certain distance, the binary representation appears similar to the continuous representation to the human eye.

This is useful for bi-level printers and displays. When viewed from a certain distance, the binary representation appears similar to the continuous representation to the human eye. There are several dithering methods, which can be split into two categories: ordered dither and error diffusion [5]. Error diffusion was used for this work as there is better contrast performance and reduced ordered artifacts compared to ordered dithering. Error diffusion uses an adaptive algorithm based on a fixed threshold to produce a binary representation of the original input. Each pixel value is modified by minimizing errors caused by the thresholding at previous pixel locations. In this way the thresholding error is diffused to adjacent pixels, hence the name of the method. The method of error diffusion is outlined in the following steps for each pixel:1. Threshold the value of the pixel of the original or modified image using a fixed threshold value.2.

The method of error diffusion is outlined in the following steps for each pixel:1. Threshold the value of the pixel of the original or modified image using a fixed threshold value.2. Calculate the absolute error between this thresholded value and the original value. 3. Diffuse this error by modifying the value of adjacent pixels in the original image using a filter.4. Repeat from step 1 until all pixels have been processed.The actual proportions of the error diffused to adjacent pixels is determined heuristically and a typical filter for this in 2D is that proposed by Floyd and Steinberg [4] which is shown in Figure 1. This filter is passed over the image during step 3 listed above where x is the current pixel.

This filter is passed over the image during step 3 listed above where x is the current pixel. The fractions of the error specified in the filter boxes are added to the original image pixel in step 4.x7/163/165/16 1/16Figure 1 – 2D filter proposed by Floyd and Steinberg for diffusing the binary thresholding error to adjacent pixels, where x is the current pixel.3D error diffusion is less common due to fewer potential applications. This is a simple extension of the 2D process where the error is diffused in 3D space to voxels around the current voxel.

This is a simple extension of the 2D process where the error is diffused in 3D space to voxels around the current voxel. Again, the proportions of the original error to diffuse are determined heuristically and the filter used in this case was that proposed by Lou and Stucki [5], which is shown in Figure 2.2/42x8/422/422/428/42 2/422/422/428/42 2/422/422/42Slice iSlice i+1Slice i+2Figure 2 – 3D filter proposed by Lou and Stucki for diffusing the binary thresholding error to adjacent voxels in the current and adjacent slices, where x is the current voxel.For this work, the above error diffusion methods were used to define lattice cells constructed from cylindrical members. This method is demonstrated first with the simple 2D cantilever plate problem shown in Figure 3a using the Floyd-Steinberg filter and is then extended to a 3D problem using the Lou-Stucki filter. The first stage is to conduct a finite element analysis (FEA) of the problem.

The first stage is to conduct a finite element analysis (FEA) of the problem. This is used to calculate the stress or some other measure of the part’s performance and the results can be plotted as an image as shown in Figure 3b. The handling of multiple problem load cases can be achieved by combining stress results from the individual analyses by taking the maximum for each pixel across the load cases. This correlates to taking the minimum grayscale level value for each pixel. This grayscale image is then subjected to an error diffusion algorithm as was explained earlier which generates the binary representation of variably spaced black dots on a white background shown in Figure 4. There are several issues that have to be addressed to make this method useful:1. How to map the stress values to the grayscale fringe image? 2. How to reduce the effect of local boundary condition stress concentrations? 3. How to provide some local control over generated node spacing, i.e.

How to reduce the effect of local boundary condition stress concentrations? 3. How to provide some local control over generated node spacing, i.e. min/max spacing? 4. How to control the quantity of generated nodes?Figure 3 – a) Cantilever plate 2D test case problem and b) grayscale stress fringe result from FEA where darker grey = higher stress.Figure 4 – Variably spaced black dotsVariably spaced black dots from error diffusion.Contrast modification to reduce skewing effect of stress concentrations at boundary Figure 5 – a) Contrast modification to reduce skewing effect of stress concentrations at boundary Contrast modification to reduce skewing effect of stress concentrations at boundary , and b) Error diffused result for resized image. condition application nodes, and b) Error diffused result for resized image. condition application nodesof these issues are related to the first issue.

condition application nodes, and b) Error diffused result for resized image. condition application nodesof these issues are related to the first issue. Regarding issue two, Regarding issue two, due to the nature of the FE method, it is usual for the degrees of freedom of individual nodes to due to the nature of the FE method, it is usual for the degrees of freedom of individual nodes to due to the nature of the FE method, it is usual for the degrees of freedom of individual nodes to be constrained and for loads to be applied to individual nodes. This results in unrealistic values be constrained and for loads to be applied to individual nodes. This results in unrealistic values be constrained and for loads to be applied to individual nodes. This results in unrealistic values s which skews a linear mapping of stress values to grayscale being calculated for these nodes which skews a linear mapping of stress values to grayscale being calculated for these node values.

This can be accommodated by modifying the mapping of the stress to grayscale values by values. This can be accommodated by modifying the mapping of the stress to grayscale values by values. This can be accommodated by modifying the mapping of the stress to grayscale values by using a simple image processing technique as shown in Figure 5a.

This can be accommodated by modifying the mapping of the stress to grayscale values by using a simple image processing technique as shown in Figure 5a. In this stage, the contrast of In this stage, the contrast of using a simple image processing technique as shown in the image is adjusted to reduce the upper limit of the stress value that is mapped to a grayscale adjusted to reduce the upper limit of the stress value that is mapped to a grayscale adjusted to reduce the upper limit of the stress value that is mapped to a grayscale value of zero (black).The second and third of these issuesIssue three is more involved and requires a quantitative link between the user is more involved and requires a quantitative link between the user is more involved and requires a quantitative link between the user requirements and the dithering method. The maximum and minimum sizes of the lattice cells can of the lattice cells can requirements and the dithering method.

The maximum and minimum sizes of the lattice cells can of the lattice cells can requirements and the dithering method. The maximum and minimum be affected by application requirement or because of manufacturing constraints. For example, in be affected by application requirement or because of manufacturing constraints. be affected by application requirement or because of manufacturing constraints. the case of the selective laser melting (SLM), the process will only currently self-support up to a the case of the selective laser melting (SLM), the process will only currently self the case of the selective laser melting (SLM), the process will only currently self It would also not be desirable for the lattice cell certain maximum overhang horizontal distance.

sizes to be too small as there are minimum feature size limitations due to the powder size and to be too small as there are minimum feature size limitations due to the powder size and laser beam diameter, and issues regarding powder removal from dense lattice regions. The laser beam diameter, and issues regarding powder removal from dense lattice regions. The laser beam diameter, and issues regarding powder removal from dense lattice regions. The resulting sizes of the lattice cells are dependent on the dithering method, specifically the size of resulting sizes of the lattice cells are dependent on the dithering method, specifically the size of resulting sizes of the lattice cells are dependent on the dithering method, specifically the size ofm overhang horizontal distance. It would also not be desirable for the latticeyscale values themselves within this image. The effect of these the input image and the grayscale values themselves within this image.

It would also not be desirable for the latticeyscale values themselves within this image. The effect of these the input image and the grayscale values themselves within this image. The effect of these parameters on the resulting lattice cell sizes was investigated the resulting lattice cell sizes was investigated. Both issue two and three relate to Both issue two and three relate to the first issue in that all are modifications to the mapping between the analysis results and the the first issue in that all are modifications to the mapping between the analysis results an the first issue in that all are modifications to the mapping between the analysis results an dithered results. This paper focuses on how to adjust the mapping to ensure certain user dithered results. This paper focuses on how to adjust the mapping to ensure certain user dithered results.

This paper focuses on how to adjust the mapping to ensure certain user dithered results. This paper focuses on how to adjust the mapping to ensure certain user dithered results. This paper focuses on how to adjust the mapping to ensure certain user requirements are satisfied.Regarding issue four, the overall quantity of the generated dithered points can be Regarding issue four, the overall quantity of the generated dithered points can be Regarding issue four, the overall quantity of the generated dithered points can be iginal image as this resolution determines the controlled by simply modifying the size of the original image as this resolution determines the controlled by simply modifying the size of the or resolution of the resulting dithered image. A reduced size dithered result in shown in Figure 5b. resolution of the resulting dithered image. A reduced size dithered result in shown in resolution of the resulting dithered image.

A reduced size dithered result in shown in Figure 5b. resolution of the resulting dithered image. A reduced size dithered result in shown in resolution of the resulting dithered image. A reduced size dithered result in shown inExperimentsSome simple test image samples were generated of varying resolution and grayscale imple test image samples were generated of varying resolution and grayscale level imple test image samples were generated of varying resolution and grayscale as shown in Figure 6. The grayscale ranges from 0 to 255, but an upper limit of 250 was used as The grayscale ranges from 0 to 255, but an upper limit of 250 was used as The grayscale ranges from 0 to 255, but an upper limit of 250 was used as In total there were 28 samples and in a pure white region, no dithered points would be present. In total there were in a pure white region, no dithered points would be present.

In total there were in a pure white region, no dithered points would be present. nt grayscale level images was subjected to the 2D error diffusion method each of these constant grayscale level images subjected to the 2D error diffusion method Steinberg filter. The resulting dithered points were connected using a using the Floyd-Steinberg filter. dithered points were connected using a elaunay triangulation meshing technique and the areas of these elements constrained Delaunay triangulation meshing technique and the areas of these elements calculated. Figure 7 shows the average element size for each sample plotted against the sample shows the average element size for each sample plotted against the sample shows the average element size for each sample plotted against the sample edge length, with fitted curves and their respective coefficients,calculated using MATLAB.

For example, what is the image size required for an average element size of 1mm2 in high stress (black) regions? This is a minimum element size requirement in the regions of grayscale value of 0 and so using the power law relationship fitted to the plotted results (equation 1) this was calculated to be approximately 70 x 70 pixels.where y is the average element area (mm2) and x is the sample edge length (pixels). So an average minimum element size can be predicted using this relationship, but this modification of the image size also affects the maximum element size. For the 70 x 70 pixel image size, the maximum average element size was calculated using equation 2 to be approximately 60mm2.Sample size: 122550100250c6 0.1301 0.1139 0.112 0.1066(1)(2)If this size is too big then the only way to achieve a desired requirement is to modify the grayscale value of this region. This can be achieved using grayscale darkening where the value of the level is changed.

This can be achieved using grayscale darkening where the value of the level is changed. The magnitude of the modification can be determined from the fitted relationships between the variables. For this example, the grayscale value required to achieve a maximum element size of 10mm2 with an image size of 70 x 70 pixels is 225. Therefore, the extent of the modification is -25. This value can be subtracted from the grayscale map to reduce the range of the levels. Because the lowest value of the range is 0, the smallest element size is unchanged. The desired extreme element sizes are therefore defined by modifying the image size for the minimum size and then modifying the grayscale level for the maximum size. Clearly, this could also be achieved in reverse. To validate that this approach translated onto a larger part, the simple test case image was modified based on the image size and darkening requirements.

To validate that this approach translated onto a larger part, the simple test case image was modified based on the image size and darkening requirements. The resulting Delaunay triangular mesh is shown in Figure 8 and the minimum and maximum regions had average element sizes of 1.08mm2 and 9.75mm2, respectively, compared with targets of 1mm2 and 10mm2, respectively. It can be seen from the mesh that there are some heavily distorted elements in the low stress regions. This is due to the meshing constraint of only connecting up the dithered points to the boundary corner points. These distortions do not affect the quality of the FEA as these triangular elements are replaced by 1D beam elements at each edge.

These distortions do not affect the quality of the FEA as these triangular elements are replaced by 1D beam elements at each edge. However, for buckling or aesthetic reasons, this constraint can be relaxed by adding additional elements to these regions which results in a more uniform mesh.Target: 1mm2 Actual: 1.08mm2Target: 10mm2 Actual: 9.75mm2Figure 8 – Good confirmation of resulting minimum and maximum element size targets based on experimental sample results.The end result of this process is a lattice structure that has smaller cells in regions of higher stress where more reinforcement is required and larger cells in regions of lower stress where less reinforcement is required. This is a spatially functionally graded lattice structure that is achieved efficiently as no iterative optimization cycles are required. To ensure that the lattice can withstand the loading requirements, the lattice is next subjected to an optimization of the lattice member dimensions.

To ensure that the lattice can withstand the loading requirements, the lattice is next subjected to an optimization of the lattice member dimensions. It is unfeasible to treat each member as a separate design variable, and ideally only a single, or a few variables would be required. The functional grading has been achieved by the dithering method so a single design variable will ensure that the lattice is strong and stiff enough. This will result in an efficient lattice design method.method can be extended to 3D with a few modifications. lattice generation method can be extended to 3D with a few modifications. from the FEA results by Firstly, volumetric analysis results are required, which can be obtained from the FEA results by Firstly, volumetric analysis results are required, which can be obtained in plane dimensions.

taking slices through the volume of the 3D part at the same resolution as the taking slices through the volume of the 3D part at the same resolution as the in plane dimensions It is important that these resolutions are the same or the lattice will be distorted in a particular It is important that these resolutions are the same or the lattice will be distorted in a particular It is important that these resolutions are the same or the lattice will be distorted in a particular diffusion filter shown in Figure 2 Secondly, as would be expected, the 3D error diffusion filter shown in dimension. Secondly, as would be expected, the 3D error . Thirdly, a constrained 3D Delaunay mesher is required that can handle non- should be used. Thirdly, a constrained 3D Delaunay mesher is required that can handle non . Thirdly, a constrained 3D Delaunay mesher is required that can handle non d connect interior and boundary dithered points.

Thirdly, a constrained 3D Delaunay mesher is required that can handle non . Thirdly, a constrained 3D Delaunay mesher is required that can handle non d connect interior and boundary dithered points. One such mesher is convex point envelopes and connect interior and boundary dithered points. One such mesher is Tetgen by Hang Si [10]. Fourthly, because the dithered points are distributed through the volume Fourthly, because the dithered points are distributed through the volume Fourthly, because the dithered points are distributed through the volume method is needed to of the component with points not necessarily on the boundary faces, a of the component with points not necessarily on the boundary faces, a method nsure mesh connectivity between lattice and solid non-design regions. This is an important ensure mesh connectivity between lattice and solid non This is an important consideration to ensure analyzability of the whole component.

This is an important ensure mesh connectivity between lattice and solid non This is an important consideration to ensure analyzability of the whole component. As shown in the flowchart in As shown in the flowchart in consideration to ensure analyzability of the Figure 9, this can be achieved by separating the boundary face and volume points. The boundary his can be achieved by separating the boundary face and volume points. The boundary his can be achieved by separating the boundary face and volume points. The boundary points can then be used to generate a surface triangulation mesh to be used as a constraint for the points can then be used to generate a surface triangulation mesh to be used as a constraint for the points can then be used to generate a surface triangulation mesh to be used as a constraint for the mesher along with the interior points. In parallel, they can be projected onto the solid non along with the interior points.

In parallel, they can be projected onto the solid non along with the interior points. In parallel, they can be projected onto the solid non-design along with the interior points. In parallel, they can be projected onto the solid non regions of the component and used as constraints for a solid mesh. In this way, the meshes, when regions of the component and used as constraints for a solid mesh. In this way, the meshes, when regions of the component and used as constraints for a solid mesh. In this way, the meshes, when can be ensured to be compatible. brought back together, can be ensured to be compatible.This whole lattice generationFlowchart showing stages required to ensure mesh connectivity between the latticed Figure 9 – Flowchart showing stages required to ensure mesh connectivity between the latticed Flowchart showing stages required to ensure mesh connectivity between the latticed region and the adjoining solid non-design regions.

region and the adjoining solid nonConclusionsThis paper has presented a novel method of generating a functionally graded lattice based This paper has presented a novel method of generating a functionally graded lattice based This paper has presented a novel method of generating a functionally graded lattice based on the results of finite element analysis. The spatial variation in the area/volume of the lattice on the results of finite element analysis. The spatial variation in the area/volume of the lattice on the results of finite element analysis. The spatial variation in the area/volume of the lattice cells is achieved using error diffusion to convert a continuous ton e image into binary form. Some cells is achieved using error diffusion to convert a continuous tone image into binary form.

Some cells is achieved using error diffusion to convert a continuous tone image into binary form. Some simple experiments were presented that allow the adjustment of the maximum and minimum simple experiments were presented that allow the adjustment of the maximum and minimum simple experiments were presented that allow the adjustment of the maximum and minimum lattice cell sizes based on determined parametric relationships . It was found that the image size based on determined parametric relationships. It was found that the image size where A where A is the lattice was related to the lattice cell size by a power relationship, was related to the lattice cell size by cell area, c1 and c2 are the coefficients and It was also found that the grayscale value was related to the cell size by a double exponential relationship, grayscale value was related to the cell size by a double exponential relationshipare the coefficients and L is the image edge length.

It was also found that tare the coefficients, and V is the grayscale value. The use of these relationships was demonstrated with a simple test case these relationships was demonstrated with a simple test case these relationships was demonstrated with a simple test case example.where A is the lattice cell area,is the lattice cell area, c3, c4, c5 and c6 are the coefficients, andThis method is efficient as it only requires a single FE analysis to be carried out per load case. It does not require iterative optimization algorithms to vary the cell size. This creates a spatial variation of the lattice which can be used as a basis for subsequent optimization of the lattice member dimensions. In some more complicated cases, it may be required that multiple FE analyses be used to improve reliability. Alternatively, an unpenalized variable density topology optimization method could be used prior to the dithering stage.

Alternatively, an unpenalized variable density topology optimization method could be used prior to the dithering stage. The current method based on a single FE analysis was demonstrated with a 2D example and the modifications required for extension to 3D explained. Further work will be on determining a parametric relationship between the stress results and greyscale value map to more closely predict lattice element sizes for particular stress levels. This will enable the method to be applied to practical problems.AcknowledgementsThe authors would like to thank the UK Technology Strategy Board (TSB) [11] forfunding this work and the Atkins project partners.References[1] Watts, D. M., A genetic algorithm based topology optimisation approach for exploiting rapid manufacturing's design freedom, PhD Thesis, Loughborough University, 2008 [2] Wang, H., and Rosen, D.

LS and ManufacturingOlli Nyrhilä, Juha Kotila, Maria Latikka, Jouni Hänninen, Tatu Syvänen, EOSFinland, Lemminkäisenkatu 36, 20520 Turku, FinlandAbstractDirect Metal Laser Sintering (DMLS) has been used for manufacturing prototypes, functional metal components and prototype tools for more than 10 years. During this period the technology has advanced to a level where direct production of complex metallic parts for various applications is everyday life and manufacturing with its various challenges is its main target. The shift from prototyping to production requires changes in the technology and also in the organizations taking part in the shift.This paper presents the latest status of the DMLS technology and materialsdevelopment trends for different application areas using EOSINT M270 laser sintering machine.

Commercially launched materials include presently biomedical materials like Titanium and Cobalt Chrome alloys, ultra high strength Maraging Steel alloy, Stainless Steels and other high-end engineering materials. In addition, there are many materials which have been developed for evaluation purposes, waiting for industrial applications.IntroductionThe first generation DMLS technology was introduced in 1994-1995. The cooperation between EOS GmbH and Electrolux Rapid Development resulted in a novel DMLS technology and the EOSINT M250 machine. This machine was equipped with 200W CO2 laser, Galvo scanners, flat field optics and suitable mechanics for taking care of powder dispensing, recoating and building platform movements [1]. This platform was used when EOS first generation materials including DirectMetal and DirectSteel families were introduced.

This platform was used when EOS first generation materials including DirectMetal and DirectSteel families were introduced. The DirectMetal materials were bronze based metal powders with various layer thicknesses and were predominantly used for prototype tooling. This material is still widely used in tooling and part production because of its easy processing and speed. The DirectSteel family consists of two steel based powder mixtures, the first being EOS “carbon Steel” and the second, DirectSteel H20, EOS first generation production grade tool steel.During the years it became obvious that the industry requires new and bettermaterials to be able to fully utilise generative manufacturing methods in either replacing existing production or developing completely new ways of manufacturing. To be able to do this EOS developed EOSINT M270.

To be able to do this EOS developed EOSINT M270. It is a new machine concept based on a solid- state, dual-focus laser and redesigned mechanics enabling faster processing with higher detail resolution than the previous generation machine. Moreover, the M 270 provides a more efficient platform for current DMLS materials and especially for ongoing development of new powders and processes for a wider spectrum of future applications.The ytterbium fibre laser used in the EOSINT M 270 has a beam quality M² ofalmost 1.0, which enables it to be focused to less than 100µm (0.0040 in.) beam diameter over the entire 250 mm x 250 mm build area. With 200 Watt power this corresponds to an average power intensity of up to 25 kW/mm². This laser also has a shorter wavelength than CO2 lasers, giving higher absorption in metals and therefore resulting in higher effective power and enabling higher build speeds.

This laser also has a shorter wavelength than CO2 lasers, giving higher absorption in metals and therefore resulting in higher effective power and enabling higher build speeds. The variable laser focus enables a very fine focus for best possible detail resolution to be combined with fast, efficient exposureof large areas by using a defocused, i.e. broader laser beam. Of course this can be utilised also in the optimization of processes for various materials [2].DMLS and manufacturingIt is now a widely accepted fact that generative methods, at least some of them, are real manufacturing methods. However, insufficient experience makes it difficult to make any general rules for the feasibility of the certain materials, batch sizes or part geometries. How well DMLS or any other generative method manages to compete technically or economically depends on the case.

How well DMLS or any other generative method manages to compete technically or economically depends on the case. A completely new design enables the use of generative advantages which increases the success rate, whereas replacing an existing production is generally much more difficult. One of the main advantages seen already in the early days of Rapid Prototyping was the reduction of restrictions related to manufacturing. This is of course true only if the part design is not an existing one and everything can be started from the beginning. The reality is different. So far only a limited amount of cases known to the authors have been of this kind. Most of the present manufacturing cases carried out with generative methods seem to be justified otherwise with the speed of the process chain or with features impossible for the conventional manufacturing methods.The advantages of the generative methods depend also on the application.

Forexample, the advantages of conformal cooling, which has been one of the most obvious new features throughout the history of the generative industry, was proven really only recently. It required further technical development of the technology, materials and design software, but above all, courage from the service providers and end users to take the necessary steps to implement the technical possibilities. The case presented later in this paper shows a very simple and good application of conformal cooling and still only scratches the surface of the advantages which will be shown in the future.The target of more or less all the parties in the generative methods business is toenter the manufacturing world. This has led to an increasing demand of reliability and quality assurance.

This has led to an increasing demand of reliability and quality assurance. The industry is far away from the early days’ stage when it was a sheer miracle if something usable came out of the machine but still more actions have to be taken before the standards of the ‘normal’ industry are fulfilled. This is a complicated task because expectations are high. It is also not fully understood that generative methods are actually part of the process industry, which means that most of the properties are created and defined in the process, inside the machine.Manufacturing Materials After the introduction of the new platform EOSINT M270 several new materials, the second generation, have been introduced. All of them are so called production grade materials, which means that existing standards can be applied in describing the properties. Of course, because the production method is new, dedicated standards do not exist yet, but gradually they will be developed.

Of course, because the production method is new, dedicated standards do not exist yet, but gradually they will be developed. Therefore normally cast or wrought references are used.The first second generation material for the M270 machine was CoCrMoW biomedical alloy for dental restorations and bridges. The material is called CobaltChrome SP1. The material is to be veneered with a ceramic material, which is done at high temperature. Therefore the coefficient of thermal expansion of this material is tuned to besuitable for this. Figure 1 shows a plateful (250x250 mm2) of dental restorations made in one job.Figure 1 200 CobaltChrome SP1 dental restorationsThe second Cobalt based superalloy is CobaltChrome MP1. This material wasdeveloped for the manufacture of biomedical implants. The properties of the alloy meet biomedical standards ASTM F75 and ISO 5832-4 of cast CoCrMo alloys and standards ASTM F1537 and ISO 5832-12 of wrought CoCrMo alloys.

The properties of the alloy meet biomedical standards ASTM F75 and ISO 5832-4 of cast CoCrMo alloys and standards ASTM F1537 and ISO 5832-12 of wrought CoCrMo alloys. Figures 2 and 3 show two different implants made of CobaltChrome MP1.Figure 2 CobaltChrome MP1 jaw implantFigure 3 CobaltChrome MP1 knee implantInterestingly, the MP1 material has also applications in engineering. Because of its high temperature properties it can be used for example in hot sections of engines and low volume run parts for jet engine development. In these applications the normal choice would be nickel based superalloys.

In these applications the normal choice would be nickel based superalloys. The availability of similar properties combined with very fast manufacturing and increased complexity offers new possibilities to the industry.In stainless steels there are currently two options, 17-4 and StainlessSteel PH1.The 17-4 is based on the metallurgy of conventional 17-4 PH material but the properties are more like 316L stainless, characterized mainly by the extremely high ductility. The demand for a material which has more conventional type properties was however so high that EOS developed a second stainless steel, this time fully precipitation hardenable material called StainlessSteel PH1. The properties of this material are equal or better than the properties of the conventionally manufactured 17-4 PH. One of the main advantages is that parts come out of the machine in solution annealed and quenched condition. Only precipitation hardening, H900 or similar simple heat treatment, is needed.

One of the main advantages is that parts come out of the machine in solution annealed and quenched condition. Only precipitation hardening, H900 or similar simple heat treatment, is needed. Figure 4 shows some properties of the PH1 material in as manufactured and after different heat treatments.

Only precipitation hardening, H900 or similar simple heat treatment, is needed. Figure 4 shows some properties of the PH1 material in as manufactured and after different heat treatments. SA stands for solution annealed, H900 and H1150 precipitation hardening in 900 and 1150 degrees Fahrenheit, respectively.UTSYield and Ultimate Tensile Strength StainlessSteel PH1138287713311267898103010201486106414361366100312621278800.00900.001000.001100.001200.001300.001400.001500.00H900, 482 ºC, 1.5 hoursSA, 1040ºC 0,5 hours, water quenchSA (1040ºC)+ H900, 482 ºC, 1 hourSA (1040ºC)+ H900, 482 ºC, 4 hoursSA (1040ºC)+ H1150, 625 ºC, 4 hoursH1150, 625 ºC, 4 hoursAs-manufacturedHeat treatment conditionStrength (MPa)YS(0.2%)Figure 4 Some properties of StainlessSteel PH1 in different conditionsThe company’s response to series production tooling requirements is MaragingSteel MS1 material. It is an ultra high strength steel material which is also suitable for part production.

It is an ultra high strength steel material which is also suitable for part production. The material was originally developed for aerospace part production but the unique combination of properties makes it extremely useful in generative processes. The parts come out of the machine as solution annealed and additionally they require only aging in moderate temperature. Compared to traditional tool steels the thermal treatment is really simple and produces only negligible dimensional changes. Figure 5 shows an example of MaragingSteel MS1 injection molding insert with cooling channels.Figure 5 MaragingSteel MS1 tool insertThe material which has the most expectations among the generative methods isTitanium. From engineering point of view it is a very attractive material but unfortunately very expensive and difficult to process, especially to meet the medical criteria.

From engineering point of view it is a very attractive material but unfortunately very expensive and difficult to process, especially to meet the medical criteria. Therefore the medical companies are putting considerable amount of effort in pushing the generative methods beyond the quality requirements for the industry.

Therefore the medical companies are putting considerable amount of effort in pushing the generative methods beyond the quality requirements for the industry. The same applies of course for aerospace industry which has also started its evaluations.EOS also has different grades available, both commercially pure and Ti6Al4V.For the medical industry Ti6Al4V ELI grade is the most challenging and this can also be produced in M270 Titanium version.Table 1 shows some main properties of five different EOS materials.CobaltChrome MP1CobaltChrome SP1StainlessSteel 17-4StainlessSteel PH1Ti6Al4VDescriptionCoCrMo based alloy for biomedical implantsCoCrMoW based alloy for dental crowns and bridges17-4 PH type stainless steel alloy for biomedical instrumentsPrecipition hardenable stainless steel for biomedical instrumentsTi6Al4V alloy for implants and engineeringRelative densityApprox. 100 %Approx. 100 %Approx. 100 %Approx. 100 %Approx.

There are several companies which have commercialised the production of CoCr bridges and copings using EOS or competing equipment. One of the first ones was a German company Sirona Dental Systems. Figure 6 shows a small and a larger bridge made by Sirona.Figure 6 CoCr bridges (Courtesy of Sirona Dental Systems)As mentioned earlier in the paper, conformal cooling now seems to be finding its applications in tooling. Figure 6 shows a simple but effective way of cooling individual locations in tool. There are 200 MaragingSteel MS1cooling pins built on top of M270 baseplate. They all have a very narrow cooling channel going from the bottom all the way to the top allowing very effective cooling.Figure 7MaragingSteel MS1 coolingpinsThe same principle can be utilised in various types of hot processes in controlling temperatures.

This has many effects on molding processes, for example shortened cycle times and also improved quality of the molded products, not to mention reduced need for tool service.ConclusionsMany of the generative manufacturing methods are gradually approaching the necessary quality levels required in different fields of manufacturing. The requirement levels are of course different depending on the application. In many industries, such as medical and aerospace, the requirements are based on legislation and therefore without fully fulfilling them the application doesn’t exist. EOS has set clear goals to itself to be a manufacturing process provider also for the most demanding industries. The selection of materials and properties described in this paper shows the progress towards the goal.References1. T. Syvänen, “Direct Laser Sintering of Metals – Material and Process Development”, 1996, M.Sc. Thesis, Tampere University of Technology, Finland.2. Syvänen, T., Heugel, M.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDMP Monitoring as a Process Optimization Tool for Direct Metal Printing (DMP) of Ti- 6Al-4VNachiketa Ray, Manisha Bisht, Lore Thijs, Jonas Van Vaerenbergh, Sam Coeck3D Systems, Grauwmeer 14, 3001 Leuven, BelgiumAbstractMetal Additive Manufacturing (AM) has evolved as a production technique for rapid prototyping as well as high volume precision manufacturing. In this work, DMP Monitoring, a new feature of 3D Systems’ direct metal printer, ProX® DMP 320 has been used as a tool for process parameter optimization. The effect of the variations of process parameters like layer thickness, laser power, scan speed and hatch spacing on the physical and mechanical properties of the additively manufactured Ti-6Al-4V samples have been investigated.

In addition to the conventional post-processing evaluation methods like Archimedes’ density, X-ray CT and tensile testing, new in-situ process monitoring tools are assessed and compared with the traditional evaluation methods.KeywordsLaser powder bed fusion, Ti-6Al-4V, process optimization, monitoring, X-ray CT1.IntroductionConventional processing of cast and wrought alloys followed by numerically controlled machining used to be an indisputable choice in the metal industry until the additive age. Additive manufacturing can be described as ‘a process of joining materials to make objects from 3D model data, usually layer upon layer as opposed to subtractive manufacturing methodologies’ [1]. The most commonly used metal additive technology on an industrial scale is powder bed fusion, where layer upon layer of powder is deposited on a build platform (of ideally the same material) and are fused together using a laser or electron beam.

Although metal additive manufacturing has revolutionized the approach towards near net shaping, rapid prototyping and high volume precision production, the technology presents itself with its own set of challenges. One of the challenges to transfer this technology from prototyping to volume production is to ensure part quality [2]. Therefore, process optimization is an integral part of the technology, since it determines the stability of the process and reinforces assurance amongst the users regarding the reproducibility of part quality.Titanium is one of the metals with high specific strength (ratio of strength to density) amongst the naturally occurring elements and combined with its excellent corrosion resistance makes it an appropriate candidate for state-of-the-art applications. Ti-6Al-4V is the most widely used Ti alloy as it accounts for more than 50% of Ti usage in the world [3,4]. It is challenging to produce Ti-6Al-4V components using conventional means.

Ti-6Al-4V is the most widely used Ti alloy as it accounts for more than 50% of Ti usage in the world [3,4]. It is challenging to produce Ti-6Al-4V components using conventional means. Castings for this alloy costs 2-3 times the cost of superalloy castings due to specific process conditions, forging is unable to produce complex shapes and powder metallurgy processes usually result in lower mechanical properties [5,6]. As a result, laser based powder bed fusion has become popular and has led to extensive research regarding the use of Ti-6Al-4V processing, especially for aerospace and biomedical applications [7–12].Process parameter optimization of Ti-6Al-4V powder bed fusion in order to achieve near theoretical density has been of considerable interest in both experimental [13,14] and nachiketa.ray@3dsystems.com2244modelling [15] techniques.

Sometimes the experimental methods come across process instabilities which lead to defects in the as-printed components like porosity (lack of fusion, key holes and gas pores), cracks and warpage in addition to the expected homogeneously distributed porosities (caused by non-optimal process parameters) [16–19]. This stochastic nature of the process has led to a lot of research in the field of in-situ process control and monitoring using various methodologies [20–24] in order to track down defect formation during the process and mitigate the risk of failure during application. However in this work, DMP Monitoring, a new feature of 3D Systems’ direct metal printer, ProX® DMP 320 has been used as a process optimization tool instead of a process control technique. Ex-situ characterization techniques like density, X-ray CT and mechanical properties are compared with the data collected by DMP Monitoring.2.

Ex-situ characterization techniques like density, X-ray CT and mechanical properties are compared with the data collected by DMP Monitoring.2. ExperimentsAtomized Ti-6Al-4V ELI (extra low interstitial grade) powder having a chemical composition which complies with ASTM F3001 and ASTM B348 (LaserForm® Ti Gr23 [25]) was additively manufactured in a commercial ProX® DMP 320 from 3D Systems. The direct metal printer (DMP) is shown in Fig 1 (a) and the three basic components of the laser powder bed fusion setup are shown in Fig 1 (b). It consists of an optical system guiding the laser beam in order to selectively melt the powder, a build envelope in combination with the powder deposition unit and the DMP Monitoring suite. The optical system is equipped with a 500 W tterbium fiber laser, having a wavelength of 1070 nm, which is guided and focused on the build envelope using galvanometer scanning mirrors and f - θ lens respectively.

The build Y envelope is 275 × 275 mm in the X-Y plane and travels in Z up to 420 mm. The Z travel after each scanned layer determines the layer thickness of the process. The powder which is stored in two feedstocks placed on the left and right of the build envelope is uniformly deposited using a scraper from both sides every alternate layer. (𝐿𝐿𝐿𝐿 [𝜇𝜇𝜇𝜇])Fig 1. (a) ProX® DMP 320 (b) Schematic illustrating the laser powder bed fusion process in a ProX® DMP 320 (c) Design of the blocks additively manufactured in this work.The DMP Monitoring suite comprises of DMP Vision and DMP Meltpool. DMP Vision is a camera used to capture images every layer post scanning and deposition. DMP Meltpool is used to monitor the light intensity emitted from the process using an off-axis Ge photodiode2245in real time, details of which can be found in our previous work [26].

DMP Meltpool is used to monitor the light intensity emitted from the process using an off-axis Ge photodiode2245in real time, details of which can be found in our previous work [26]. An example of the signal acquired by DMP Monitoring during the direct metal printing process is shown in Fig 2 where it synchronizes the data collected after every layer for aided visualization.Fig 2. DMP Monitoring user interface (UI) showing DMP Meltpool (a) and DMP Vision (b) acquisition in real time for a given layer. The DMP Meltpool bitmap is a false color image with the color representative of the observed intensity.of 30µm and 60µm were chosen to print the blocks having dimensions as shown in Fig 1 (c). Blocks of each layer thickness were printed with varying laser power having a constant hatch spacing.

Blocks of each layer thickness were printed with varying laser power having a constant hatch spacing . The variations in laser power (90% - 140% of the standard) and scan speed 𝑠𝑠 (75% - 250% of the standard) were made accordingly to have an increasing amount of (𝑣𝑣 volumetric energy density , which can be defined as the average energy imparted per unit volume on the powder bed [27–29] as shown in Eq. (1).In this work, two layer thicknessesand scan velocities (𝐿𝐿𝐿𝐿 [𝜇𝜇𝜇𝜇])−1(𝑃𝑃 [𝑊𝑊 ])[𝜇𝜇𝜇𝜇 · 𝑠𝑠])𝑠𝑠 (𝐻𝐻[𝜇𝜇𝜇𝜇])(𝑉𝑉𝑉𝑉𝑉𝑉)6−3𝑃𝑃 𝑠𝑠 𝐿𝐿𝐿𝐿 · 𝑣𝑣The density of all the blocks were measured using Archimedes’ principle. Two samples from each layer thickness were chosen to be scanned with X-ray computed tomography (CT) in a Phoenix Nanotom (GE) with a voxel size of approximately 15µm to calculate the amount of porosity in the blocks.

VGSTUDIO MAX (Volume Graphics) was used for data analysis.𝑉𝑉𝑉𝑉𝑉𝑉 =𝑠𝑠 × 10[𝐽𝐽 · 𝜇𝜇𝜇𝜇]𝐻𝐻The DMP Meltpool data was acquired by the detector and logged as a function of time and the data analysis was carried out separately and is not a part of the collection software (Fig 2). The data analysis as a function of time is non-trivial and also impractical. On the contrary, mapping the signal intensity as a function of spatial coordinates on the powder bed is much more comprehensible. The X and Y coordinates of the laser position at the powder bed level were logged every 20 μs along with its corresponding total emitted intensity acquired over those 20 µs.

The X and Y coordinates of the laser position at the powder bed level were logged every 20 μs along with its corresponding total emitted intensity acquired over those 20 µs. The DMP Meltpool data (which contains the X, Y, Z co-ordinates and the corresponding intensity recorded by the photo diode) for each layer was stored in a separate file and was analyzed using an algorithm which detects the intensity fluctuations above a certain threshold and over a certain spatial range and tags them as DMP Meltpool events. The2246(1)criteria for DMP Meltpool events detection were based on holistic rules established in our previous study [26]. The probability of a DMP Meltpool event resulting into a defect is higher if it manifests itself in more than one layer in the vicinity of the same XY co-ordinate. The volume of each DMP Meltpool event , was added up for all the events following the above criteria was calculated using Eq. (2).

The volume of each DMP Meltpool event , was added up for all the events following the above criteria was calculated using Eq. (2). Each Ti-6Al-4V block having a volume,, and the ‘DMP Meltpool event to part volume ratio’𝑖𝑖was analyzed individually in this work. 𝑚𝑚 )(𝑣𝑣)(𝑛𝑛)(𝜌𝜌𝑉𝑉𝑛𝑛 𝑖𝑖=1 ∑ 𝑣𝑣𝑖𝑖Near net shaped round tensile test coupons were printed in the direction perpendicular to the build platform with the same process parameters as the four blocks selected for X-ray CT analysis. No heat treatment or surface finishing was carried out for the tensile coupons prior to testing. For each process parameter set, six samples having dimensions as recommended by ASTM E8M [30] with a gauge diameter of 4mm were printed. Room temperature tensile tests were performed according to ASTM E8M on the as-built samples with a strain rate of 0.005 min-1 in the elastic regime and 0.4 min-1 in the plastic regime.𝑚𝑚𝜌𝜌=× 100%𝑉𝑉3.

Results and discussionAfter the direct metal printing process, the blocks were detached from the Ti build platform, and the density of each block was measured using Archimedes’ principle and is shown in Fig 3 as a function of volumetric energy density (a, b) and scan velocity (c, d) for both layer thicknesses. The theoretical density was taken as 4.42 g·cm-3 and it can be seen that the relative density never goes above 99.3%. This can be attributed to the measurement method [31] and in order to verify that, the pixel density of one of the samples with the highest Archimedes’ density was measured on a grinded and polished cross-section which corresponded to a value of 99.97% for an Archimedes density of 99.35%. It should be noted that the relative density increases with increasing VED, but above a certain value, depending on the layer thickness, the relative density achieves a stable value.

It should be noted that the relative density increases with increasing VED, but above a certain value, depending on the layer thickness, the relative density achieves a stable value. The initial low value of density can be attributed to porosity caused by lack of fusion as a result of insufficient overlap between the melt pools. Further increasing the VED does not reflect any drastic changes in density measured by Archimedes’ principle for LT30 (Fig 3 (a)), but a small drop in density is observed for LT60 (Fig 3 (b)) which can be associated with the formation of keyhole porosities.for the blocks are also plotted as a function of VED in Fig 3 (a, b). For the layer thickness 30µm, no trend was observed, however for layer thickness 60µm the does not correspond to the volume fraction of porosity, but acts as a mere indication of the volume fraction of pores.

𝑚𝑚 𝑚𝑚 However, it’s interesting to note that although the relative density remains constant for LT60 𝜌𝜌 𝜌𝜌 above VED of 30 J·mm-3, the value decreases up to 40 J·mm-3 and then becomes stable. This would be an indication that at least VED of 40 J·mm-3 is required to achieve near theoretical density. As mentioned before, the current analysis requires an abrupt fluctuation in the DMP Meltpool signal of adjacent vectors which persists for consecutive layers to flag an event, which is not expected to be observed in the signature of a keyhole. In order to compare the volume fraction of events detected by DMP Meltpool with the actual porosity fraction, X- ray CT analysis was performed on two samples with high Archimedes’ density from each layer.

The measured porosity fraction for the four samples are thickness, one with low and high and the CT pore indicated in Fig 3 (a, b) and its corresponding to part volume ratio appear to follow a similar trend but the absolute values should not be 𝑚𝑚 compared as mentioned before.The calculated𝑚𝑚decreases with increasing VED. It should be noted that the𝜌𝜌𝑚𝑚𝜌𝜌is highlighted. Both the𝑚𝑚𝜌𝜌𝑚𝑚𝜌𝜌𝜌𝜌2247(2)are plotted as a function of increasing scan velocity. Multiple data points for a given velocity value are due to differences in laser power. The range of the X axes in both Fig 3 (c) and (d) are kept identical in order to provide some reference to the reader. Again for LT30, no clear trend is observed, while for LT60 the increases with increase in scan velocity.

Again for LT30, no clear trend is observed, while for LT60 the increases with increase in scan velocity. This provides an indication that if a process parameter set is chosen just based on the Archimedes’ density, parameters with higher scan velocity have a significant chance of having more events than parameter sets with lower scan 𝜌𝜌 velocity. However, it should be taken into account that all the data points shown in Fig 3 (d) have different values of VED. In order to have a more accurate perspective, two interesting VED ranges for the layer thickness of 60µm have been considered and their corresponding relative density and have been plotted against the scan velocity in Fig 4. For the sake of comparison, the scale has been kept identical to Fig 3 (d). It can be perceived that for high VED ~40 J·mm-3, which also corresponds to lower scan velocities, no trend in can be observed.

It can be perceived that for high VED ~40 J·mm-3, which also corresponds to lower scan velocities, no trend in can be observed. However, for low VED ~30 J·mm-3, corresponding to higher scan velocities, increases with increasing scan velocities. This again leads to the fact that DMP Meltpool is not 𝑚𝑚 only suitable for detecting lack of fusion porosities caused by low VED (predictable) but also defects which form randomly.In Fig 3 (c, d) both the relative density and𝑚𝑚𝜌𝜌𝑚𝑚𝑚𝑚𝜌𝜌𝑚𝑚𝜌𝜌Fig 3. Relative theoretical density measured using Archimedes’ method and its corresponding event to part volume ratio (%) is shown as a function of the volumetric energy density applied during the powder bed fusion process for layer thickness 30µm (a) and 60µm (b). The ratio of the pore to part volume (%) measured by CT is also indicated in (a) and (b) and its corresponding event to part volume is highlighted.

The ratio of the pore to part volume (%) measured by CT is also indicated in (a) and (b) and its corresponding event to part volume is highlighted. (c) and (d) illustrate the relative density and event to part volume ratio (%) as a function of increasing scan velocity for layer thickness 30µm and 60µm respectively.2248𝜌𝜌Fig 4. Relative density and event to part volume ratio (%) as a function of increasing scan velocity for layer thickness 60µm plotted in the same scale as Fig 3 (d) for VED range (a) 38- 42 J·mm-3 and (b) 28-32 J·mm-3.Room temperature tensile testing was carried on the as-built coupons printed with the same process parameters as the four blocks selected for X-ray CT analysis.

The yield strength (YS), ultimate tensile strength (UTS) and the plastic strain were calculated from the stress- strain curves and the average and 2σ (twice the standard deviation) of six tests for the four process parameters are summarized in Fig 5.1500YSUTSPlastic strainLT30LT601300) a P M1100(h t g n e r tS90070050030042472840Volumetric energy density (J·mm-3)Fig 5. Room temperature tensile properties of as-built samples printed in the direction perpendicular to the build platform for layer thickness 30µm and 60µm, having the same laser parameters as the samples highlighted in Fig 3 (a) and (b). The four process parameters also correspond to the blocks selected for X-ray CT analysis.It is well known from literature that the microstructure of as-printed, non-heat treated ) as a result of the high cooling rates Ti-6Al-4V consists of fine acicular martensite ( characteristic of the process [12,27].

As a result of which direct metal printed Ti-6Al-4V tend to have high yield and tensile strength but relatively low ductility [32–34]. Small changes in porosity are directly reflected in the ductility of the as-printed material [17]. The plastic strain observed in Fig 5 for a given layer thickness and VED is consistent with the CT pore to part as shown in Fig 3 (a) and (b). With increasing porosity or DMP Meltpool volume ratio and events, the plastic strain decreases. It is important to note that a higher VED may not correspond′𝛼𝛼𝜌𝜌𝑚𝑚2249121086420)%(n a r t sic i t s a Plto lower porosity levels. Although the relative density measured using Archimedes’ principle for LT30 is comparable for VED of 42 and 47 J·mm-3, the plastic strain, CT density and DMP Meltpool density were all lower for the samples printed with the higher VED (47 J·mm-3).

It was also observed that samples with higher CT pore to part volume ratio and had a large scatter in the plastic strain data for both layer thicknesses. Therefore, solely from 𝑚𝑚 this work it is the intention of the authors to adopt as an indication or possibility of defect formation rather than a measure of porosity since further verification and validation of the DMP Meltpool data analysis is required.𝑚𝑚(1 − 𝜌𝜌)𝑚𝑚𝜌𝜌4. ConclusionsIn this work, several blocks were additively manufactured by laser powder bed fusion for layer thickness (LT) of 30µm and 60µm with varying laser power and scan velocity for a constant hatch spacing. DMP Monitoring was used as an in-situ process monitoring tool to identify non-conformities during the process. Analysis of the DMP Meltpool data was compared with X-ray CT measurements performed on selected samples with similar relative density measured using Archimedes’ principle.

Analysis of the DMP Meltpool data was compared with X-ray CT measurements performed on selected samples with similar relative density measured using Archimedes’ principle. A strong correlation was observed between the DMP Meltpool event to part volume ratio and the CT pore to part volume ratio. For LT30 no trend was observed between the volumetric energy density (VED) and , however for LT60, a correlation was observed for lower VED and higher scan velocities. In order to ) compare with the mechanical properties, room temperature tensile tests were performed on as- built near net shape samples printed perpendicular to the build platform. The plastic strain . It was shown that showed a strong correlation with the CT pore to part volume ratio and the VED and relative density measured using Archimedes’ principle were not always a good indication for selecting process parameters required to print defect free components.

main Decomposition and Space Filling Curves in Toolpath Planning and GenerationBertoldi M., Yardimci M. A., Pistor C. M., Giiveri S. I. University of Illinois at ChicagoABSTRACTEver increasing computer performance, along with significant developments in CADI CAM technology and high precision digital motion controllers lead to rapid and significant developments in the field of rapid prototypingjfabrication. Together, these elements offer a wide range of possible approaches in the toolpath planning issue; two main sets have been analyzed and applied to Fused Deposition Modeling process.Domain decomposition is a frequently used technique in computational methods. Within the context of present study, this approach is used to divide arbitrary layer geometries into smaller regions of simpler shape. A foreseeable advantage in such an approach is maximizing strength characteristics through thermal management.

A foreseeable advantage in such an approach is maximizing strength characteristics through thermal management. This is achieved by utilizing space filling curves which are mathematical entities that offer the possibility of building a wide range of structures, covering the surface of a single layer with one continuous curve. To evaluate the proposed concepts, ABS structures and ceramic green bodies have been successfully built.OVERVIEWThe term toolpath planning is defined here as the collection of operations performed toprovide the Rapid Prototyping hardware with the set of building commands.Although computer-controlled machines are now commonly available, the algorithms for driving them automatically are still subject of research. In CNC machining for example, human intervention and ingenuity are still necessary to design and build the necessary jigs for clamping of the workpiece.

In CNC machining for example, human intervention and ingenuity are still necessary to design and build the necessary jigs for clamping of the workpiece. The integration of RP equipment with PC-based controllers offers the possi bility of using common software development tools and resources for toolpath planning.Fused Deposition hardware is very similar to basic operation of pen-plotters and the machine level controllers are. based on .the same language. As a •. result, they .can. work more efficiently with patterns (i.e.

based on .the same language. As a •. result, they .can. work more efficiently with patterns (i.e. porous structures) that have "polylines" as the basic geometric features, compared to the current mode of operation, which works on filled polygons (monolithic models)..267The algorithms that have been developed and applied to the Fused Deposition can beconsidered in two main categories:..Slice level manipulationDomain decomposition is a common technique in computational methods, used here to divide arbitrary layer geometries into smaller regions of simpler shape... Machine level manipulation.Space filling curves are mathematical entities that offer the possibility of building a wide range of structures, covering the surface of a single layer with one continuous curve.SLICE LEVEL MANIPULATIONIn the current study domain decomposition was demonstrated using a Stratasys 1650 modeler.

The standard Stratasys software {QuickSlice™) performs the translation from the 3D CAD model to the 2YlD model through a slicing algorithm. At this point, the geometry of the part is described with a finite set of 2D layers, in order to match the SFF building paradigm. For certain applications, it can be more useful to directly generate or modify the geometry of the desiredstructure at slice level, i.e. in modeling of three dimensional objects from medical images where the original data is a set of 2D images. Furthermore, working on 2D contours allows the use of simpler and more common algorithms.The slice information is contained in a SSL file, where it is described as a contour repre senting a closed planar surface. In the standard software processing, the slices are subsequently filled with a set of parallel lines, representing the roads.

In the standard software processing, the slices are subsequently filled with a set of parallel lines, representing the roads. At the SSL level, the slice represents the building unit and it belongs to a different "group" to which the following properties can be assigned:.. Material .. Build style .. Contour (yes/no) .. Road width .. Air gap between roads .. Raster orientation, which is usually set to [+45 -45] for two consecutive levels.The main constraint is the fixed nature of the raster, which is always made of straight parallel roads. This affects the physical structure of the part and all its characteristics, including mechanical properties, void contents, surface and bonding quality and accuracy.

This affects the physical structure of the part and all its characteristics, including mechanical properties, void contents, surface and bonding quality and accuracy. Furthermore, the topology of the parts that can be created is limited to standard solid structures.A set of routines and programs has been developed in a combined FORTRAN/MATLABenvironment, in order to explore the possibilities of the process.DOMAIN DECOMPOSITIONDomain decomposition techniques are mostly used in computational analysis, when the mathematical model describing the physical phenomena cannot be solved in closed form. The idea is to decompose the 2 or 3-dimensional domain into regions of simpler geometry (usually polygons in 2D and polyhedra in 3D) and to discretize the equations on the obtained grid, in order to generate an approximate solution.

This approach, which had its real boost with the advent of computers, has requested the development of more refined and accurate mesh genera tor algorithms, especially for unstructured grids.In Rapid Prototyping, the physical domain can be assimilated to the part to be built in the 3D representation, where the computational domain is the solid model (STL file). However, most RPISFF techniques are based on a layered manufacturing paradigm and the eventual domain decomposition of the STL model cannot be reproduced in the physical part. Instead of consider ing the entire solid model, the domains to be decomposed are the single layers and the final mesh is a new layer, with the same extemalcontour but subdivided in a variable size and shape mesh.With this approach, it is possible to increase the number and manipulate the character istics of the domains, which at this level represent the building units.

This results in a more flexible and customizable building process.DELAUNAY TRIANGULATIONIn two dimensions, a triangulation of a set V of vertices is a set T of triangles whose vertices collectively form V, whose interiors do not intersect each other, and whose union completely fills the convex hull of V.One of the most common techniques for generating triangular grids is based on Delaunay triangulation. The Delaunay triangulation D of V, introduced by its author in 1934, is the graph defined as follows. Any circle in the plane is said to be empty if it contains no vertex of V in its interior (Vertices are permitted on the circle). Let u and v be any two vertices of V. The edge UV is in D if and only if there exists an empty circle that passes through U and V.

Let u and v be any two vertices of V. The edge UV is in D if and only if there exists an empty circle that passes through U and V. An edge satisfying this property is said to be Delaunay.Although initial Delaunay triangulation algorithms have been developed for simply connected convex polygons [1], [2], current algorithms [3], [4], are able to generate variable density Delaunay triangulations within multiply connected arbitrary geometry polygons; some of them find applications in numerical modeling problems [5]. Delaunay triangulation is perfectly suited for a domain described with a set of segments; the program trian.exe developed in 1996 by Shewchuk [6] is· able to generate triangles at a rate of 105 points in 30 seconds.

The structure of the algorithm is presented here in pseudo-code:in our case,for i=l:number_of_layersfor j=l:number_of_slices269read slice file translate current slice -> input_domain call trian.exe read output file write new sliceendendThefunctio~ssl.mhas beellimplemented in MATLAB. The old set properties from the inputSSL file generated with Qu.ickSliceare read, and the file is scanned for number of layers and number of distinct. domains within' each of them. The structure of the layer is then .read, translated and written into a file understandable by trian.exe. Once the domain decomposition has beenperfonned, the generated triangular mesh is read, and each triangle is assigned to one of the previously set group (each of them can have different properties,as described above).

Finally, theInew data structure is written in SSL format and can be reentered in the standard software for toolpath generation.In the tested cases, two different command...line option were used with trian.exe:p, which causes the program to generate a Planar Straight Line Graph (PSLG), which is a collection of points and segments.

Segments are edges whose endpoints are points in the PSLG, and whose presence in any mesh generated from the PSLG is enforced.a x, which constrains the maximum area of a single triangle to x.Each layer of the tensile test bars depicted in Figure 1 have been decomposed with a dif ferent maximum area constraint; the resulting triangles have been randomly assigned to a group with variable road orientation [0 ±30 ±45 ±60 90].DalaunaytrtangLJaDon- b_iso.sslDalaunaytrtangLJation - b_iso.ssl~~~~~~20>'102040608010012014016020406080100120140180DelaunaytrtangLJation - bJsosslDalaunaytrtangLJatlOn - b_isossl2020>-10~102040608010012014016020406080100120Figure 1. Domain decomposition of a tensile test barAnother example is depicted in Figure 2.

Domain decomposition of a tensile test barAnother example is depicted in Figure 2. The interior of the octagon was tessellated with a constrained Delaunay triangulation, and constructed triangles were defined as individual con tours within the SSL file. Road width was fixed to 0.5 mm; individual triangles were then ran domly assigned to a set of group with variable raster angles and air-gaps (0.3 to 0.7 mm)?7n2. Domain decomposition structureMACHINEGENERATIONThe only way to fully explore the possibilities of the FDM process is to directly generate the machine level language, in order to avoid all the constraints of the standard toolpath proces sing software. It is in fact possible to skip the solid model generation and slicing steps (STL SSL files) and directly output the final machine file (SML).tests is driven by an Asymtek A-201 digital motion controller.

The A-20t· controls the x-y movement of the deposit ing head, the z movement of the stage, and the rotation of the two electrical servo-motors mount ed on the head that feed the thermoplastic wire. into the two liquefiers. The controller uses Autolllove Control Language (ACL) for programming [7]; Stratasys has implemented a slightly modified version of this. language, called Stratasys Machine Language (SML). It is similar to Hewlett Packard's PCL used to control plotters and all commands are strings of ASCII text.The StratasysFDM 1650 machine used for the experimentalThe structure of the SML file has been reverse-engineered in order to separate the com mon control sequences from the real flow. and motion control commands.

and motion control commands. A set of routines has been developed in FORTRAN and then ported to MATLAB; the structure of the basic algorithm with a detailed description of each function is shown here in pseudo-code:smlhead () purge ()for i=l:number_of_layers_basebase () (curvebrk)endfor i=l:number_of_layerslevel () purge () curve () toolpath () (curvebrk) endcurve ()end smltail()271where:smlheadpurge level curvetoolpathcurvebrksmltailEstablishes communication parameters, initializes the motion controllers and prepares the FD machine for the build. Performs on-the-fly cleaning operations (tip-wipe). Switches materials. ifnecessaryandpositions the z-stage for the current layer. Sets the flow-control parameters for the current road Generates motion controlcol11mands for the deposition of the road Writes a pausing se~uence.TheI11.otion controller has a very limited on-board memory that can store only 256 consecutive commands.

When the buffer is full, the transmission is interrupted and the stored commands executed. Communication link is reactivated near the end of 256-command set and the next batch is loaded to the memory. Performs final cleaning operations, clears the memory and puts the machine in stand-by mode for part removaLIt should be noticed that the 3D topology of the final structure can still be maintained; infact, each layer can be generated with a different toolpath.FRACTAL SPACE FILLING CURVESA space-filling curve (SFC) is a continuous line that can cover a region of space without intersecting itself. Applications of space filling curves can be found in digital images processing, ray tracing optimization and fluid dynamics. The idea of using fractals in RP is due to the in trinsic nature of the curve: it can cover each layer without starting-stopping sequences.

The idea of using fractals in RP is due to the in trinsic nature of the curve: it can cover each layer without starting-stopping sequences. The length of the unit steps within SFC's can be set to a value equal to the width of the roads, in order to completely fill each layer.The Hilbert space-filling curve (Figure 3) has been found particularly useful and conse quently applied to the toolpath planning issue. The developed MATLAB program, shown in pseudo-code, generates the Hilbert curve, which is stored in a vector (toolpath), translated into SML language and written to file. The function main () contains the connectivity matrix (a [4x4x3] array) that defines the topology of the curve and recursively calls the hilbc () function:main ()conn_matrix = [ ... J hilbc (order);end (main)function hilbc (order)if (order>O)for n=l:orderhilbc (order-I); generate_toolpath ();endend?7?Figure 3.

J hilbc (order);end (main)function hilbc (order)if (order>O)for n=l:orderhilbc (order-I); generate_toolpath ();endend?7?Figure 3. Hilbert curve generation .One of the main advantages of this approach is related to the thermal history of the part produced; from microscopy images, it has been observed that whenever the head deposits long straight lines, the bonding quality between two contiguous roads can be poor or non-existent. This is due to the thermal expansion and subsequent cooling and shrinking of the single road. With the fractal approach, it is possible to depose one layer reducing significantly the length of the movements of the head. In this way, the void contents can be lowered and the bonding quality can be improved.

In this way, the void contents can be lowered and the bonding quality can be improved. The main disadvantage is the build time, which is 2-4 times longer compared to the standard toolpath, due to the exponentially increased number of motion commands.The microstructure generated by 0/-90° stacking of a 4th order Hilbert curve is shown inFigure 4-a; Figure 4-b shows the part produced using this approach (ABS material).0/-90 degree stacking of4th order Hilbert SFC2000 .---~-~-~----,1500'"1000500500100015002000Figure 4. Microstructure toolpath (a) and real part (b)APPLICATIONSPossible applications of the proposed techniques cover all the Fused Deposition uses, where custom physical and mechanical properties of the parts must be achieved. The software routines developed can be adapted to different needs and are easily configurable and customiz able.273In particular, one interesting application concerns the mechanical behavior of the FDM parts.

The standard. raster settings cover every slice with a set of parallel roads, causing a strongly Qrthotropic. behavior of the. obtained. part. Decomposing the domain and assigning the properties to .each new contour independently allows more control on the internal microstructure. S.peci111ensbuilt •.••. with .fractal space filling curves were tested, showing better mechanical propettiescompared to the standard parts.Another interesting application taking direct advantage of the proposed techniques is inporous structure generation algorithms; see [8] for more details.CONCLUSIONSWith the new programming tools available, it is possible to generate different routil}es for developing new toolpathplanning techniques. Some of them have been presented, discussed and deemed effective in the toolpath planning issue, either as a stand-alone product or interfaced with the standard processing software.

Some of them have been presented, discussed and deemed effective in the toolpath planning issue, either as a stand-alone product or interfaced with the standard processing software. Future work will include integration with other domain de composition algorithms and STL three-dimensional domain decomposition.REFERENCES1. 2.3.4.5.6.7. 8.Bowyer A., "Computing Dirichlet tessellations," The Computer Journal, vol. 24 no. 2, pp. 162-166, 1981. Watson D. F., "Computing the n-dimensional Delaunay tesselation with application to Voronoi polytopes," The Computer Journal, vol. 24, no. 2, pp. 167-171, 1981. Guibas L. 1., Knuth D. E., Sharir M., "Randomized incremental construction of Delaunay and VoronoI diagrams," Algorithmica, vol. 7, no. 4, pp. 381-413, 1992. Shimada K.,Physically-Based Mesh Generation: Automated Triangulation of Surfaces and Volumes via Bubble Packing, Ph.D. thesis, Massachusetts Institute of Technology, 1993 M. A. K. Posenau, D. M.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperDreaming of Data: Examining Data Augmentation for Machine Learning in Additive Manufacturing Glen Williams, 1 Martha Baldwin,2 Timothy W. Simpson,1 Nicholas A. Meisel,1 Christopher McComb21School of Engineering Design and Innovation, The Pennsylvania State University, University Park, PA 16802 2Department of Mechanical Engineering, Carnegie Mellon University, Pittsburgh, PA 15213Abstract The data generated during additive manufacturing (AM) practice can be used to train machine learning (ML) tools to reduce defects, optimize mechanical properties, or increase efficiency. In addition to the size of the repository, emerging research shows that other characteristics of the data also impact suitability of the data for AM-ML application.

In addition to the size of the repository, emerging research shows that other characteristics of the data also impact suitability of the data for AM-ML application. What should be done in cases for which the data in too small, too homogeneous, or otherwise insufficient? Data augmentation techniques present a solution, offering automated methods for increasing the quality of data. However, many of these techniques were developed for machine vision tasks, and hence their suitability for AM data has not been verified. In this study, several data augmentation techniques are applied to synthetic design repositories to characterize if and to what degree they enhance their performance as ML training sets. We discuss the comparative advantage of these data augmentation techniques across several canonical AM-ML tasks.1.

We discuss the comparative advantage of these data augmentation techniques across several canonical AM-ML tasks.1. IntroductionStudying design repositories through metamodels can provide useful insight into whether or not an engineering design dataset will be successful in training a deep learning construct. However, what should be done if a dataset is estimated to be insufficient? Instead of simply giving up, researchers presented with incomplete, small, or otherwise low-quality datasets often turn to data augmentation to enhance their dataset prior to training a machine learning model. In this study, several data augmentation techniques are applied to artificial design repositories to characterize if and to what degree they enhance their performance as machine learning training sets. The intersection of additive manufacturing (AM) and machine learning (ML) presents numerous, complementary opportunities for design engineers.

The intersection of additive manufacturing (AM) and machine learning (ML) presents numerous, complementary opportunities for design engineers. Manufacturing products using AM allows designers to take advantage of opportunistic AM attributes, such as part consolidation to reduce the number of parts in an assembly by increasing the complexity of individual parts [1] and hierarchical complexity to make parts that are lighter while remaining strong [2]. These beneficial attributes of AM can help designers focus more on problem solving than on limitations of traditional, tooling-driven manufacturing.In the same way that AM enables designers to manufacture increasingly complex and general shapes that might better solve a problem, ML could allow them to better predict how well those diverse solutions will perform in a multitude of ways.

Some examples of researchers using ML to analyze AM parts include predicting AM build metrics [3], predicting thermal properties of AM designs during a build [4], classify parts as manufacturable or not [5], and prediction of product1642geometric values to result in desired stress-strain performance [6]. However, both AM and ML are not perfect solutions, and many open questions regarding how to best utilize these technologies remain. ML approaches are generally most successful when developers have access to large training and validation datasets. Unfortunately, engineering data can be costly to create, obtain, and manage, resulting in datasets that are much smaller or more poorly labeled than those in some other fields.

Unfortunately, engineering data can be costly to create, obtain, and manage, resulting in datasets that are much smaller or more poorly labeled than those in some other fields. This problem is only compounded by the fact that AM technologies are relatively new and heuristics that might be useful for labeling datasets are the topic of contemporary research [7,8] and therefore in flux.Data augmentation can be a powerful tool to enhance the performance of ML constructs while reducing their tendency to overfit [9,10]. It should be considered as an additional standard practice as part of the ML development process. Many ML studies focus on making incremental improvements to ML architectures while making fundamental assumptions that often hold a training and test dataset constant. In this type of study, an architecture-focused ML development process is often followed (see Figure 1).

In this type of study, an architecture-focused ML development process is often followed (see Figure 1). In this architecture-focused ML approach, the researcher has selected a) a dataset to use and hold constant and b) an architecture to test and present as an incremental improvement if successful. In the ML feature optimization loop shown in Figure 1, automated training algorithms, such as backpropagation, are used to iteratively alter ML construct weights and biases to improve the performance of the construct.Figure 1: An architecture-focused ML development process iterates only on the ML construct architecture.While this architecture-focused approach is time-tested for fields with highly standardized, agreed- upon benchmark datasets, it may not as applicable for other fields. For instance, the space at the intersection of Design for AM (DfAM) and ML, which we refer to together as DfAM-ML, suffers from a lack clearly and readily available datasets.

For instance, the space at the intersection of Design for AM (DfAM) and ML, which we refer to together as DfAM-ML, suffers from a lack clearly and readily available datasets. For instance, surrogate modeling studies may require the generation of a new dataset for specific problems. The collective lack of understanding surrounding the impact of design repository attributes and data augmentation techniques on ML outcomes is a gap that could be further explored.

The collective lack of understanding surrounding the impact of design repository attributes and data augmentation techniques on ML outcomes is a gap that could be further explored. To investigate this gap, this study presents a broader ML development process, the augmented dataset nested iterative ML development process (see Figure 2).1643Figure 2: In the augmented dataset nested iterative ML development process, the model features, ML architecture, and training and evaluation datasets may all be varied to achieve greatest performance.In the augmented dataset nested iterative ML development process, the same ML feature optimization is conducted, but additional loops to augment the dataset are conducted as well. Although these additional iterative loops may increase the research and development labor cost in the short term, they ideally will result in a greater likelihood of success of the entire ML development process.

This study investigates specific practical techniques that could be applied to the new dataset augmentation loop in Figure 2.In this work we investigate the details of how data augmentation can be useful in improving sub- optimal datasets in the context of DfAM-ML development. By studying a variety of data augmentation techniques across many, hierarchically complex design repositories with differing levels of constituent and aggregate complexity, and in the context of a variety of ML classification and regression problem types we aim to advance the heuristics around the DfAM-ML development process to increase the speed and efficiency with which useful ML constructs may be designed, trained, and tested. Specifically, we study the following research questions:1. Which data augmentation techniques are most effective in improving the performance of ML training activities?2.

Specifically, we study the following research questions:1. Which data augmentation techniques are most effective in improving the performance of ML training activities?2. Do the data augmentation techniques affect trained ML construct performance similarly across different problem types?The remainder of the work is organized as follows: First, we document the methodology of our candidate data augmentation techniques (section 2). We then present our results (section 3), discussion of the results (section 4), and conclusions regarding potential future work related to this study (section 5).16442. MethodologyThis section documents the data augmentation techniques investigated (section 2.1), process for creating the artificial design repositories (ADRs) (section 2.2), overview of the CNN architecture (section 2.3), and training procedures (section 2.4).2.1.

Data Augmentation Technique Treatments In this study, five data augmentation approaches were individually applied as treatments to Artificial Design Repositories (ADRs). An ADR is a synthetic dataset which contains constituent designs that were not necessarily designed for the primary purpose of manufacturing [3]. These treatments aim to increase the performance of a design repository as a machine learning training dataset. The treatments include a) no data augmentation (control), b) duplication, c) experience, d) single randomized orientation, e) multiple orientations, and f) gaussian noise (see Figure 3).Figure 3: Overview of the data augmentation techniques investigated in this study.Each treatment from the figure is explained in further detail as follows:In the no data augmentation (control) treatment, the ADRs were included as-is with no modifications to the training dataset. See section 2.2 to better understand the generation of the ADRs.

See section 2.2 to better understand the generation of the ADRs. Including a control treatment is important to establish a baseline for comparison of the magnitude of positive or negative effects each data augmentation treatment had. In these control repositories the designs were modeled in the same orientation, resulting in a consistent placement and orientation in the voxel-based input to the CNNs. In the duplication data augmentation treatment, the existing constituent designs in the training dataset were simply duplicated once. This is the simplest way to double the size of the dataset. Heuristically, this method may be susceptible to overfitting, but is evaluated to test that hypothesis empirically. In the experience data augmentation treatment, the dataset is doubled in size by including random designs from other artificial design repositories.

In the experience data augmentation treatment, the dataset is doubled in size by including random designs from other artificial design repositories. This treatment simulates the action of augmenting a dataset the hard way: by waiting for more data to come in, gaining1645more experience. In this analogy, random constituent designs added to the ADR are akin the new experience of designing more parts and storing their data in the repository. In the single randomized orientation data augmentation treatment, the dataset size was not doubled, but instead the models were rotated to a random orientation. This treatment aims to provide the CNN a greater variety of voxel-based features during the training process than the consistent orientation designs of the control treatment.

This treatment aims to provide the CNN a greater variety of voxel-based features during the training process than the consistent orientation designs of the control treatment. In the multiple orientations treatment, the dataset size was doubled with the addition of another random orientation of the same model, instead of its replacement (as is the case in the single randomized orientation treatment). This resulted in a larger dataset that also included a greater variety of voxel-based features. In the gaussian noise treatment, normally distributed pseudo-random noise was applied to the constituent design voxel-based inputs. This gaussian noise has the effect of altering the volumetric occupancy binary voxels, which are usually only values of 1.0 or 0.0, to be decimal values between 1.0 and 0.0. The magnitude of the noise was centered around 0.05 with a standard deviation of 0.025.

The magnitude of the noise was centered around 0.05 with a standard deviation of 0.025. Each calculated random noise sample for a given voxel was added to that voxel’s value if it previously had a value 0.0 and subtracted from its value if it previously had a value of 1.0.2.2. Artificial Design Repositories The ADRs used in this study were created using procedural modeling of geometries based on parametric design templates. Parametric design templates are advantageous for ADRs because they allow for rapid automated creation of constituent designs while achieving a desired level of geometric and topological variability in those designs. This variability is what makes the ADR suitably similar to real-world design repositories for the purposes of the current theoretically- oriented study. A multistep process was used to model the constituent designs and assemble them into ADRs suitable for the current work (see Figure 4).

A multistep process was used to model the constituent designs and assemble them into ADRs suitable for the current work (see Figure 4). Specifically, the ADRs used in this study included 310 constituent parametrically designed parts. The parts were analyzed for classification and scalar regression labels relevant to DfAM-ML problems. A detailed example of the hierarchical modeling modelling process is shown in Figure 5.Figure 4: Flowchart of the overall Artificial Design Repository generation methodology16463Figure 5: Detailed illustration of the hierarchically complex modeling process demonstrating high degrees of variability of geometric and topological complexity were achieved in the artificial design repository modeling process.Creating a constituent design involves first selecting a macrostructure design template from 18 options and specifying arguments for its dimensional parameter values (see step 1 in Figure 5).

These 18 unique macrostructure design templates are constructed to have between 1 (the sphere) and 7 (the I-beam) numeric dimensional parameters. They were selected to relate to common material stock forms, such as prisms, cylinders, and beams. After a design macrostructure was selected and defined, a design could optionally have a single additive or subtractive feature applied to it through constructive solid geometry Boolean CAD operations [11] (see option 2: step 3 in Figure 5). The 5 optional features included uniform shell hollowing, circular hole, circular post, rectangular hole, and rectangular post (see option 2: step 3 in Figure 5). Parameters of features included values such as “wall thickness” and “hole depth” (see option 2: step 4 in Figure 5). Features add to the number of parameters already available when procedurally generating models, ranging from 1 added parameter (the shell) to 6 added parameters (the rectangular post).

See option 2: step 3 in Figure 5 for a diagram of the possible features.Instead of a feature, a design could optionally have a mesostructure applied to it (see option 3 in Figure 5). Although mesostructures in DfAM applications are often 3D, 2D mesostructures were used in this study because of the limits imposed by the computational demand of the sheer number of designs modeled. These mesostructures included 2D patterns of circular cutouts and 2D pattern1647rectangular prism cutouts. Applying mesostructures has a similar effect of increasing the geometric and topological complexities of designs as features, except to a higher degree.

Applying mesostructures has a similar effect of increasing the geometric and topological complexities of designs as features, except to a higher degree. The mesostructure also add parameters to the design template, in this case related to size and 2D spacing distances of the mesostructure unit cells.2.3 Convolutional Neural Networks This study employed a similar convolutional neural network (CNN) to the prior metamodel study done by Williams et al., but with fewer channels in the convolutional layers and more neurons in the fully-connected layers (see Figure 6) [12]. The input was a 64 x 64 x 64 volumetric grid, a “voxel-based model”. The voxels were binary numbers represented as 32-bit floating point values, with a 1.0 representing material presence and 0.0 representing material absence. These CNNs were trained with binary cross entropy loss function [13]. The hidden layers of the CNN were organized into an initial convolutional section fully-connected secondary section.

These CNNs were trained with binary cross entropy loss function [13]. The hidden layers of the CNN were organized into an initial convolutional section fully-connected secondary section. The layers included three alternating 3D convolutional-max-pool layer pairs, a flattening layer, and 3 fully-connected layers. The rectified linear activation function was used for all hidden layers. Each CNN model had between 33,614 trainable parameters.andoutput The dimensionality and loss function depending on the problem type studied. Binary classification networks AM manufacturability classification) used a single binary output neuron (32-bit, 1.0 or 0.0 floating point) with a sigmoid activation. These CNNs were trained with binary cross entropy [13]. Ternary function classification CNNs (Hierarchical complexity classification) used 3 binary output neurons (32-bit, 1.0 or 0.0 floating point), which labels that were one-hot encoded [14].

Ternary function classification CNNs (Hierarchical complexity classification) used 3 binary output neurons (32-bit, 1.0 or 0.0 floating point), which labels that were one-hot encoded [14]. Multi-class classification CNNs (design template classification) used 18 binary output neurons (32-bit, 1.0 or 0.0 floating point), which labels that were also one- hot encoded. Both ternary and multi-class classification CNNs were trained with categorical cross entropy loss function [15] and using a SoftMax activation function [15]. All scalar prediction CNNs (part mass, build time, minimum build time) had a single 32-bit floating point output neuronnetworksdifferedin(Go-no-goLPBFlossFigure 6: Diagram of the convolutional neural network architecture studied.1648with a linear activation function and were trained with a mean squared error loss function. In line with the learnings of Williams et al.

In line with the learnings of Williams et al. relating to the standardization of orientation in a design repository, all designs were rotated to a random orientation before their input voxel-based model was prepared to maximize CNN performance [16].2.4 Training The training procedure was as follows: First, an ADR was loaded. Next, the ADR was split into randomized 75% training and 25% testing dataset. The CNN model was then trained using the training data. The history of each trained CNNs’ loss, validation loss (either cross entropy or mean squared error), accuracy (either classification or mean squared error), and validation accuracy for each epoch were recorded. The trained CNN was then used to make predictions on the evaluation dataset, which was a constant dataset that included five times the number of constituent designs as an ADR, equal to 1,550 designs.

The trained CNN was then used to make predictions on the evaluation dataset, which was a constant dataset that included five times the number of constituent designs as an ADR, equal to 1,550 designs. This evaluation dataset consistent of a uniform random sampling of all the models and was intended to simulate analysis of future designs that would not be guaranteed to be similar to the training dataset. The CNNs were developed and trained using TensorFlow version 2 [17], Python version 3.10.4, and the “Adam” Optimizer [18]. All training was completed on an NVIDIA P6000 Graphics Processing Unit.In this study, the CNN accuracy metric of interest was either classification accuracy or mean squared error when predicting values on an evaluation dataset containing a large number of random designs external to the training ADR. Each design repository in this study was used to train 36 CNNs (1 for each data augmentation treatment for each CNN problem type).

Each design repository in this study was used to train 36 CNNs (1 for each data augmentation treatment for each CNN problem type). This replicates the activity of evaluating the data augmentation techniques 189. In total, 6,804 CNNs were trained in this study. The performance of data augmentation treatments on an individual ADR was analyzed by calculating the difference between the control CNN performance on the evaluation dataset and the augmented CNN performance. These difference values were min-max normalized prior to analysis. The classification CNNs used the classification accuracy as the performance metric and the scalar prediction CNNs used the mean squared error as the performance metric. Mean squared error was subtracted from 1.0 after normalization so that all deltas of the performance metrics across different CNN problem types could be compared on a scale in which 1.0 is the best possible value.3.

ResultsApplying the data augmentation treatments to design repositories resulted in variable increases, and sometimes decreases in trained CNN accuracy. The central tendency, spread, and distribution of data augmentation performance deltas (the difference between the augmented dataset performance and the control) were different for each data augmentation treatment in the context of each CNN problem type. Table 1 includes a summary of the mean and standard deviation of the normalized trained CNN accuracy delta for the different data augmentation treatments when applied to classification CNNs. Table 2 includes a summary of the mean and standard deviation of the normalized trained CNN accuracy delta for the scalar prediction (regression) CNN problem types. Figure 7 provides a set of plots which communicate this information visually.

Figure 7 provides a set of plots which communicate this information visually. For all the plots, the vertical values have been min-max normalized based on the range of accuracy values observed for all CNNs in that problem type.1649Table 1: The mean and standard deviation trained CNN performance delta for all data augmentation treatments applied to each classification CNN problem type.

In plots (a), (b), and (c), accuracy (the vertical axis) is the classification accuracy for the entire evaluation dataset. In plots (d), (e), and (f), accuracy (the vertical axis) is 1.0 minus the mean squared error for the entire evaluation dataset.16514. DiscussionOverall, the experience data augmentation treatment, which simulated the addition of new designs from outside the design repository over time, was clearly a high performing data augmentation technique. Experience exhibited the largest trained CNN accuracy deltas compared to the control on average for classification problems and the second largest for scalar prediction problems (see Table 1). This result confirms the intuition that new data is generally better than techniques to modify and augment data. However, the result that experience was not the best for both problem types is interesting. For classification, including more examples increased CNN performance the most.

However, the result that experience was not the best for both problem types is interesting. For classification, including more examples increased CNN performance the most. This makes intuitive sense, because in a classification dataset the frequency and balance of the output categories in the dataset is generally important [19]. However, in the scalar prediction problems, categorical labels are not relevant to the optimization of the CNN, and variability in the high-dimensional, voxel-based input through the addition of noise appears to have a particularly beneficial effect.Data augmentation techniques that increased the size of the dataset generally performed better than those which did not. For instance, the single random orientation treatment, in which the constituent designs were rotated to a random angle, but the original design orientation was not included in the training dataset, was among the lowest performing for both classification and regression problems.

The single random orientation treatment decreased the performance of the CNNs on average for some classification and regression problems. This result further supports the results from Williams et al., in which random rotations were found to be beneficial to CNN performance [16]. Although random rotations do frequently increase the performance of a neural network in this study, they do not appear to do so consistently across all problem types and when averaged over many different, hierarchically complex design repositories.

Adding more design repositories, greater complexity and variability to those design repositories, and more CNN problem types further contextualizes the impact of random rotations on trained CNN performance, showing it is not as effective as some of the other data augmentation techniques considered.The duplication data augmentation treatment outperformed the multiple orientation data augmentation treatment in both classification and regression problems, despite both treatments doubling the size of the input training dataset. This result is interesting because it suggests that the size of the input dataset, although important, is not the entire story in terms of the impact of data augmentation. This observation further supports the conclusions of the metamodel study by Williams et al., in which size was found to be the most important metadata attribute of a design repository by far [12].

This observation further supports the conclusions of the metamodel study by Williams et al., in which size was found to be the most important metadata attribute of a design repository by far [12]. The results of the current work show that the precise details of how the size of a design repository is increased can have complex effects on the overall performance of the trained CNN. In other words, a greater quantity of augmented designs appears to be often, but not always, better for CNN training outcome. Specifically, duplication might be more effective than including multiple orientations in this case because it simulates the effect of increasing the number of training epochs, essentially giving the CNN more time to train on the dataset.

Including hyperparameter variation in the experimental methodology could provide evidence to causally link the differences observed between duplication and multiple orientations.In terms of CNN problem type, classification CNNs that were trained to predict the macrostructure design template had their performances increased the most on average by the data augmentation1652techniques. CNNs trained to predict minimum build time were the most improved by data augmentation among the scalar prediction CNN problem types. Conversely, classification CNNs that were trained to predict the ternary hierarchical complexity level of a design were the least improved overall when data augmentation techniques were applied to their training datasets. In the ternary hierarchical classification problem, the greatest frequency of average decreases in CNN performance due to a data augmentation treatment was observed.

In the ternary hierarchical classification problem, the greatest frequency of average decreases in CNN performance due to a data augmentation treatment was observed. The experience and single randomized orientation treatments both decreased performance on average and the average classification delta was below zero. Part mass predicting CNNs were the scalar prediction CNNs that were least improved by the data augmentation techniques, experience an average decrease in performance from the single random orientation. These lower performing augmentation results shows that although the data augmentation techniques used in this study most often improved CNN performance, that is a possibility that they could have a negative effect. Keeping this result in mind when preparing a large dataset for a DfAM-ML development project could be crucial to best utilizing limited computational resources, as more data augmentation may actually cause a decrease in performances in some situations.

The risk of hindering DfAM-ML performance with detrimental data augmentation, as evidenced by this result, motivates further study into the general topic of data augmentation across a variety of DfAM-ML problem types.5. ConclusionIn this study, a characterization and analysis of five data augmentation techniques to improve DfAM-ML training performance of artificial design repositories (ADRs) was conducted. The magnitude of the change of performance of all design repositories after data augmentation in the context of the 6 different CNN problem types studied were analyzed.

The magnitude of the change of performance of all design repositories after data augmentation in the context of the 6 different CNN problem types studied were analyzed. In terms of our first research question, which data augmentation techniques are most effective in improving the performance of ML training activities the experience data augmentation treatment, in which entirely new designs were obtained over time, was the best performing treatment for classification problems, while addition of gaussian noise to the voxel-based input was the best performing for scalar prediction problems. Experience was consistently an effective data augmentation treatment for both classification and regression problems.

Experience was consistently an effective data augmentation treatment for both classification and regression problems. In terms of our second research question, do the data augmentation techniques affect trained ML construct performance similarly across different problem types, we found that the central tendency of performance improvement was similar in scale for different problem types, but that the distribution was highly dependent on the problem type.Overall, we conclude that, based on our analysis, the best way to augment data is to prioritize obtaining entirely new data if possible. However, given that data augmentation is typically used in situations when obtaining more data is difficult or impossible, including many orientations of the same design in the training dataset outperforms other techniques tested.

Given the large degree of variability in the performance deltas observed across different data augmentation techniques, future work should further investigate these questions. In particular, this study shows that some data augmentation techniques can cause an average decrease in trained ML performance, and thus additional study is needed to uncover additional heuristics that enable DfAM-ML developers to apply data augmentation strategically to avoid decreasing performance.1653The results of this study have implications in the field of DfAM-ML development. All the data augmentation techniques demonstrated were effective at increasing machine learning training performance overall. This result means that industrial organizations can extract greater value from their design repository by processing it prior to using it as training data. However, the study also showed that data augmentation did not always increase ML performance, and sometimes reduced the training dataset quality.

However, the study also showed that data augmentation did not always increase ML performance, and sometimes reduced the training dataset quality. This result means that AM-ML developers should not blindly attempt data augmentation but should rather incorporate it as a part of a broader strategy. Using such a strategy means training multiple ML constructs using multiple data augmentation approaches and statistically comparing the accuracies. The degree to which data augmentation techniques are iterated should be budgeted for and balanced with the computational resources available, since both data augmentation and further ML training iteration both incur greater computational cost.Future work should expand both the breadth and depth of DfAM-ML details. The breadth could be expanded by probing whether changing the ML architecture and input dimensionality affects the data augmentation performance enhancement.

The breadth could be expanded by probing whether changing the ML architecture and input dimensionality affects the data augmentation performance enhancement. The depth could be expanded by investigating finer details of the ML training process, such as when overfitting occurs during long training activities with many epochs for different data augmentation techniques. Additionally, the interactions between multiple data augmentation techniques could be further studied to understand how they may be best used in combination.AcknowledgementsThis material is based upon work supported by the National Science Foundation (Grant No. CMMI-1825535). Any opinions, findings, conclusions, or recommendations expressed in this paper are those of the authors and do not necessarily reflect the views of the sponsors.References[1] Oh, Y., Zhou, C., and Behdad, S., 2018, “Part Decomposition and Assembly-Based (Re) Design for Additive Manufacturing: A Review,” Addit Manuf, 22(April), pp.

Dry Powder Material Delivery Device for Multiple Material Additive Manufacturing S. Chianrabutra, B.G. Mellor, S. YangEngineering Sciences Unit, Faculty of Engineering and the Environment, University of Southampton, Southampton SO17 1BJ, United KingdomAbstractThis research is to develop a novel material delivery device for a next generation additive manufacturing system which is capable of directly manufacturing objects by depositing several materials layer by layer. The successful deposition of multiple material layers by using this novel dry powder printing technique reveals its great potential as a means of incorporating multiple materials in the traditional additive manufacturing process since this technology is suitable for a wide range of materials and it has the capability to manufacture 2D layers composed of multiple materials.

This paper will outline the basics of the dry powder printing technology and present and discuss selected experimental studies from our research.IntroductionCurrently, most commercial additive manufacturing systems are designed to produce parts from a single material [1]. Multiple Materials Additive Manufacturing (MMAM) using direct inkjet printing technology was pioneered by Evans and co-workers nearly 20 years ago [2] and is now researched by various groups. Objet has commercialized direct inkjet 3D printers and claim that they represent the only MMAM system. However, Objet can only print photopolymer materials, which have limited mechanical properties and other functionalities, compared to direct metal and/or ceramic printing. MMAM is a new technology that can fabricate three-dimensional multiple material (heterogeneous) objects. This technology can create multiple material objects and can vary material compositions within the layer [3].

This technology can create multiple material objects and can vary material compositions within the layer [3]. The ability to print multiple materials from an additive manufacturing system can improve either the mechanical properties of the parts or provide additional functions to the 3D printed parts [4]. MMAM technology has the potential to become an important manufacturing resource for the next generation of AM technology. This is because single material AM systems cannot fulfil the requirements of some applications that require multiple material objects from one machine, such as compliant mechanisms, embedded components, 3D circuits, human tissues, medical compatible implants etc. Moreover, the next generation of AM technology should provide full functionality, offer changeable material systems, and give the entire bonding system at an affordable price [5-7].Multiple material objects are more interesting and will be highly important in many industrial applications.

Many research institutes and companies have been developing AM technologies to allow more materials to be used in single material AM technology, and to improve the properties of the AM part and to enhance the capability of the basic process [4]. Powder based materials are very important as they can provide a wide variety of material options. The Dry Powder Printing (DPP) technique is one of the promising techniquesavailable to dispense fine powders [8]. Among the DPP techniques, the ultrasonic dispensing method has many significant advantages for depositing fine powders uniformly and controllability and has the ability to handle a wide variety of powder materials. In most of the current powder based AM systems, powders are spread using a roller onto a powder bed which made it impossible to fill multiple materials without cross-contamination [9, 10].

In most of the current powder based AM systems, powders are spread using a roller onto a powder bed which made it impossible to fill multiple materials without cross-contamination [9, 10]. Therefore, it is essential to develop processing technologies that can handle fine powder particles for use in material delivery devices for the next generation of AM systems. However, processing and handling powder materials is very challenging due to their unique properties. To deposit such materials it is necessary to understand thoroughly the powder flow in the dispensing nozzle.Dry Powder Printing (DPP) or dry powder micro-feeding or fine powder dispensing is based on vibration, resulting from the ultrasonic vibration of a piezoelectric transducer, being used to assist flow. Over the recent past this DPP method has proved capable of handling fine powders.

Over the recent past this DPP method has proved capable of handling fine powders. Pegna, who is a pioneer in this field, studied the possibility of multiple material deposition by creating a single layer of Portland cement [11] and spherical glass beads [12]. Santosa et al. demonstrated the influence of orifice diameters and under 100 µm particle sizes on flow behaviour under gravity through a hopper-nozzle [13]. Matsusaka et al. [14] investigated the microfeeding of fine powders in a capillary tube vibrated by a 20 kHz piezoelectric transducer. Takano and Tomikawa [15] developed feeding devices based on the excitation of a progressive wave in an ultrasonic transmission line. Li et al. [16] used an ultrasonic-based micropowder-feeding mechanism to form thin patterns of dry powders on a substrate which were subsequently sintered by a laser beam. Kumar et al.

Li et al. [16] used an ultrasonic-based micropowder-feeding mechanism to form thin patterns of dry powders on a substrate which were subsequently sintered by a laser beam. Kumar et al. examined the concept of multiple dry powder deposition under gravity flow including low gas pressure- assisted flow and vibration-assisted flow and developed a model to predict the flow rate under gravity of the experimental powders [17]. Jiang et al. evaluated the flowability of nanoparticle powders [18] and developed a measurement system of powder flowability based on a vibrating capillary [19]. Yang and Evans [20, 21] presented the factors and mechanisms that were responsible for the initiation and cessation of flow in a vibration controlled dispensing system. Additionally, they identified the powder characteristics controlling powder handling so as to find effective metering and dispensing methods for shape and composition control for solid freeform fabrication [22].

Furthermore, they investigated the effects of acoustic frequency, amplitude, tube diameter, mechanical damping and particle size distribution on particle deposition from open capillaries subject to acoustic vibration [23]. They invented a discontinuous dispensing device using short pulses of ultrasonic vibration to dispense the dry powders in a drop on demand format [24]. Lu et al.[8, 25-27] studied dispensing mechanisms, drop size control, dose uniformity and different design of dispensing nozzles.However, only a few studies have been carried out using ultrasonic vibration to dispense dry powders for MMAM [4]. The detailed mechanism of DPP has not been fully clarified and is needed for the technique to be a commercial success. The processing of fine powder in the micrometre size range has proved difficult [28, 29]. This is due to the feature size being driven by the size of the dispenser orifice, a smaller feature size resulting from asmaller diameter orifice.

This is due to the feature size being driven by the size of the dispenser orifice, a smaller feature size resulting from asmaller diameter orifice. For the production of drops from sub-micrometre diameter size fine powders, the process is difficult to control reliably or sometimes powders cannot be dispensed because of their poor flowability. In this work, we have systematically investigated the effects of printing parameters on printing results which reveals the basic characteristics of our powder printing device. Furthermore, our study shows that dry powder printing, driven by ultrasonic vibration, can be exploited for fine powder printing of many types of powder materials.Materials and methodsMaterialsIn this study, the selected powder materials were copper (Osprey Metals, Neath, Wales), solder (Sn63Pb37, IPS (Suzhou) New Materials Co.

Ltd., China), 316L stainless steel (Osprey Metals, Neath, Wales), tungsten carbide (Sandvik, Conventry, UK), alumina (Al2O3, BA Chemicals Ltd., Buckinghamshire, UK), CoCr (Concept Laser GmbH, Lichtenfels, Germany), 420S45 stainless steel (Osprey Metals, Neath, Wales), glass bead (Whitehouse Scientific Ltd., UK), and Glass-filled Nylon (DuraForm® GF, 3D Systems Corp., USA) in the size range 14-72 µm and their details are shown in Figure 1 and Table 1. These materials were selected to cover metals, polymers and ceramics with different densities, particle sizes and particle shapes used in AM systems.100 µm316L SS100 µmSiC100 µmCoCr100 µmCu100 µm100 µm420S45 SS100 µmWC100 µmAl2O3100 µmGlass-filled PA100 µmFigure 1 Scanning Electron Micrographs of the experimental powders.The powders were analyzed to confirm that the size range was within specification by using a Malvern Mastersizer 2000 particle size analyzer.

The particles were also characterized by a LEO 1455 VP Scanning Electron Microscope (SEM).Table 1 Physical characteristics of the experimental materials.Powder D50 (µm) Particle density(kg/m3)Cu148940SiC543220Al2O3523970Glass412300CoCr188290SnPb358400316L SS327890420S45 SS 207740WC3515500GlassSnPbGlass- filled PA 721490Experimental setupThe experimental setup consists of a computer, an analogue waveform generator (NI 6733 DAQmx card, National Instruments Corporation Ltd. Berkshire, UK), a power amplifier (50w, Sonic Systems Ltd, Somerset, UK), a glass nozzle (Pasteur glass pipette) attached to a piezoelectric ceramic ring (SPZT8-100-50x20, MPI Co., Switzerland) by an adhesive epoxy (Araldite Rapid Syringe-Epoxy Extra Strong, Huntsman Corp., USA), Z column and X-Y table (Parker Hannifin, supplied by Micromech, Braintree, UK).

The experimental setup is as shown schematically in Figure 2.Z-columnSignal VoltageGlass TubePiezoelectric RingPower AmplifierComputerSubstrateX-Y TableMotion ControllerFigure 2 Schematic diagram of the dry powder printing system.To capture images of the powder as it discharges in the experiment, a high- speed Photron Fastcam SA-1 (Photron Limited, Japan) camera attached to a c-mount adapter on a Leica Monozoom 7 (Leica Microsystem Inc., USA) macro lens was installed close to the dispenser.Experimental conditionsThe experiments were carried out under room conditions (23-28 °C and 36-45 % RH). As contaminants in the powder can reduce their mass flow rate from the dispenser and even clog the orifice a clean environment and sieving the powder before testing are essential. In this study, powders, which were kept in air tight packages, were sieved through a 100- micrometre sieve size before filling the dispenser, i.e.

In this study, powders, which were kept in air tight packages, were sieved through a 100- micrometre sieve size before filling the dispenser, i.e. all particles dispensed passed through a 100-micrometre sieve size.Results and discussionsDispensing behaviourThe powder is dispensed by a vibration-assisted system using a glass nozzle as a funnel and a computer control system. Powder drops are discharged directly from the dispenser. Initially, particles form a stable dome structure across the orifice as shown in Figure 3.

Powder drops are discharged directly from the dispenser. Initially, particles form a stable dome structure across the orifice as shown in Figure 3. Powders are dispensed through the orifice by breaking the dome structure by activating a voltage signal pulse to the piezoelectric transducer attached to the dispenser.Stable dome structureFigure 3 Dome structure of a fine powder inside the glass nozzle.During dispensing, the vibration from the piezoelectric transducer transmits energy through the glass tube to particles around the dome structure and the result is to break the dome structure and so achieve flow of the powders. On switching off the vibrations, particle- particle and particle-wall friction lead to the formation of domes causing powder flow to arrest in the nozzle.

On switching off the vibrations, particle- particle and particle-wall friction lead to the formation of domes causing powder flow to arrest in the nozzle. Figure 4 shows the sequential images captured by a high speed camera at 0.01 second increments from the start of dispensing (B) to the cessation of dispensing (G) when the powder forms a new stable dome structure (H-J) inside the glass nozzle.(A)(B)(C)(D)(E)Stable Dome StructureONONONON(F)(G)(H)(I)(J)ONOFFOFFOFFStable Dome StructureFigure 4 Sequence of images captured from the start of dispensing to the cessation of dispensing of a fine powder from the glass nozzle.Printable materialsThe ten experimental powders detailed in Table 1 were successfully discharged as can be seen in Figure 5. All of these results came from the same dispenser geometry, i.e. a 250 µm nozzle diameter with a nozzle angle of 75°.

All of these results came from the same dispenser geometry, i.e. a 250 µm nozzle diameter with a nozzle angle of 75°. Throughout this study, the experiments were carried out at an applied signal voltage of 2 Volts, standoff distance of 200 µm and moving2 mm2 mmspeed of 10 mm/s. Notably, the appropriate control of parameters for each experimental powder can improve markedly the deposited powder patterns.316L SS2 mmSiC2 mmAl2O32 mmGlass2 mmCu2 mmGlass-filled PA2 mm420S45 SS2 mmSnPb2 mmFigure 5 Single material patterns.The quality of the patterns, which is characterised in term of continuity, consistency and fineness of the edge of the printing line produced mainly depends on the powder’s flowability and the orifice size.

Based on the particle’s density and mean particle diameter all the experimental powders, except for copper and glass-filled nylon, fit into groups A and B of Geldart’s classification presented in Figure 6, The copper powder and the glass-filled nylon are located in group C. Group A powders are ideal for conveying in sliding bed flow, Group B powders are easy to fluidize and they rapidly de-aerate while Group C powders are difficult to fluidize and are cohesive [30, 31]. Moreover, all of the experimental powders in Figure 1 do not agglomerate and have discrete particles.

Moreover, all of the experimental powders in Figure 1 do not agglomerate and have discrete particles. It may be concluded that in all cases, the powders used in this study are free-flowing materials or not extremely cohesive.100,000WC10,000CuCoCr SnPbD Spoutable)3m / g k (g ρ - s ρ420S45 SS316L SSAl2O3SiCGlassB Sand-like1,000A AeratableC CohesiveGlass-filled PA10010,000 Figure 6 Experimental powders shown in Geldart’s classification (ρs-solid density, ρg-gas density, dp- mean particle diameter)101001,000dp (µm)CoCrWCPrinting of fine track widthThe printing patterns are produced by CAD programs such as SolidWorks. The printing parameters of signal voltage, standoff distance and moving speed are programmable input via software. High precision stages are used to achieve x, y and z movement.

The printing parameters of signal voltage, standoff distance and moving speed are programmable input via software. High precision stages are used to achieve x, y and z movement. The experimental studies found that high resolution depends on powder flowability, nozzle orifice geometry (nozzle diameter and nozzle angle), standoff distance and moving speed (the relative velocity between the substrate and the nozzle) as well as the signal voltage used to activate the piezoelectric transducer. Different parameters in the dispenser system need to be carefully adjusted for optimum results as the different powders have their own flow properties.

Different parameters in the dispenser system need to be carefully adjusted for optimum results as the different powders have their own flow properties. A different adjustment of parameters can result in the feature sizes shown in Figure 7.(A)(B)200 µmFigure 7 Patterns obtained from two printing conditions with the same nozzle diameter (70 µm) and the same signal voltage (2 Volts): (A) the standoff distance is 475µm, and the moving speed is 10 mm/s, (B) the standoff distance is 125 µm and the moving speed is 15 mm/s.At present, the DPP printing device using an orifice diameter of 60 µm can realize a feature size down to 85 µm as shown in Figure 8.

During the printing process the standoff distance of the deposition nozzle from the substrate was 150 µm, the moving speed was 5 mm/s and the signal voltage was 2 Volts.200 µmFigure 8 Samples with a track width of less than 100 µm.Nozzle diameterMass flow rate as a function of dispensing nozzle diameter was determined for the copper powder and solder powder of 14 µm and 35 µm mean particle size respectively. The nozzle diameters used were 110, 200, 250, 280, 320, 375 and 400 µm. The signal voltage was fixed at 2 Volts and the nozzle angle was 66°. Figure 9 shows the results of the mass flow rate for the two selected powders tested with this range of nozzle diameters. The mass flow rates increase with nozzle diameter. The results show that with the smallest nozzle diameters, the mass flow rates of the experimental powders are similar but they rapidly increase with increasing nozzle diameter.

The results show that with the smallest nozzle diameters, the mass flow rates of the experimental powders are similar but they rapidly increase with increasing nozzle diameter. Throughout the range of these nozzle diameters, the solder powder has a higher flow rate (higher power law exponent) than the copper. This reflects the fact that a free-flowing powder such as the solder powder has higher flowability and thus greater mass flow rate than a cohesive powder such as the copper powder.30y = 1.16E-07x3.21 R² = 0.997) s / g m(20e t a R w o l F s s a M10y = 8.79E-06x2.27 R² = 0.997Cu (D50=14 µm)SnPb (D50=35 µm)00100200300400500Nozzle Diameter (µm)Figure 9 Mass flow rate as a function of nozzle diameter for copper and solder powder.Signal voltageThe piezoelectric vibration enables powder printing to be achieved.

The mass flow rate can be regulated by the vibration energy generated by the signal voltage, which controls the deformation amplitude of the piezoelectric transducer. Generally, when a greater voltage is applied the vibration amplitude increases. In our study a free-flowing powder, such as the solder powder, discharged as a spray of particles as shown in Figure 10 and the spray angle increased with increasing signal voltage. For a signal voltage of 0.15 Volts, the spray angle was less than 10º while for the signal voltage of 0.5 Volts and 2.0 Volts, the spray angle was around 30º and 90º respectively. The reason might be that the high vibration energy transmitted to the powder particles results in greater vibration of powder particles around the nozzle wall. As the particles exit the nozzle they hit the edge of the orifice and thus spread out. This effect is less when using a larger nozzle diameter or using a cohesive powder such as copper.

As the particles exit the nozzle they hit the edge of the orifice and thus spread out. This effect is less when using a larger nozzle diameter or using a cohesive powder such as copper. Thus it can be seen that by choosing a suitable signal more uniform and consistent powder dispensing can be achieved and therefore more precise patterns produced.0.15 Volts0.5 Volts2.0 VoltsFigure 10 Sequence of images obtained from a high speed camera at 0.05 second increments showing the successive stages of discharge of solder powders. The nozzle diameter is 250 µm.Standoff distanceStandoff distance is the distance of the tip of the nozzle to the top of the substrate. From the previous discussion, it can be seen that powders normally spread out, the spray angle depending on the signal voltage. Therefore, the standoff distance should be as small as possible.

From the previous discussion, it can be seen that powders normally spread out, the spray angle depending on the signal voltage. Therefore, the standoff distance should be as small as possible. However, the lower limit to standoff distance depends on the moving speed as the standoff distance must be sufficient to avoid powder blocking flow between the tip of the nozzle and the top of the substrate. Figure 11 shows the results of printing at different standoff distances. The track line at the 125-µm standoff distance results in the tip of the nozzle coming into contact with the powder resulting in compressing the printed line and might interrupt the powder flow at the outlet.Standoff distance125 µm375 µm625 µm1 mm/sd e e p sg n i v o M5 mm/s10 mm/sFigure 11 Sequence of images obtained from a high speed camera for copper powder dispensing with different moving speeds and standoff distances.

The nozzle diameter is 250 µm.Moving speedMoving speed is the relative speed between the nozzle and the substrate. A fast moving speed can increase the printing speed of the process. However, the moving speed must not be so fast as to produce a discontinuity in the printed line. Figure 12 shows that copper printing with moving speeds of 20 mm/s and 25 mm/s cannot create complete printed lines. It indicates that the maximum moving speed in this test should be less than 20 mm/s.25 mm/s20 mm/s15 mm/s10 mm/s5 mm/sFigure 12 Printing results for copper powder obtained from different moving speeds. The nozzle diameter is 250 µm, the signal voltage is 2 Volts and the standoff distant is 200 µm.Multiple material patternsFigure 13 demonstrates the capability of printing multiple powder materials. Three sets of optimal parameter were used to produce a fine pattern. The pattern combines stainless steel, tungsten carbide and copper powder.

Three sets of optimal parameter were used to produce a fine pattern. The pattern combines stainless steel, tungsten carbide and copper powder. A system consisting of three dispensers was installed on the z column and the x-y table was synchronised by the motor controller.Stainless SteelTungsten CarbideCopper1 mmFigure 13 Multiple material pattern.AdvantagesThis DPP technique can print various materials and little material preparation needs to be carried out. Using separate nozzles to deposit powder material in selective areas is an accurate, efficient and easy method which avoids contamination between materials. The device has few components resulting in less maintenance and an economical cost. Different materials can be printed simply by adding a new dispenser to the system. Utilising a nozzle based method, powder materials can be delivered between and within layers to make multi- oriented interfaces in a final part.

Utilising a nozzle based method, powder materials can be delivered between and within layers to make multi- oriented interfaces in a final part. It is believed that using a dry powder dispensing system integrated with current powder-based additive manufacturing systems, such as laser sintering, laser melting or binder jetting, would increase their ability to produce multiple material parts.LimitationsA limitation of this device is that the process requires a small standoff distance to provide high resolution and to avoid spreading due to the spray effect. Additionally, accidental vibration from the environment, e.g. movement of the x-y table, can affect the controllability of powder dispensing. However, this can be overcome by mounting the device in a system with smooth movement and adequate damping.

movement of the x-y table, can affect the controllability of powder dispensing. However, this can be overcome by mounting the device in a system with smooth movement and adequate damping. Furthermore, the dispenser must be installed vertically because the device combines vibration and gravitational force to discharge material.ConclusionsIn this paper, a novel Dry Powder Printing (DPP), based on an ultrasonic dispensing device where powder flow can be regulated by controlling input to the piezoelectric transducer, is introduced. The development of this novel DPP method would allow the layering of high quality multiple material patterns. By programming the print heads, this powder printing technique can achieved selective area deposition of different dry powder materials using ultrasonic activation without sophisticated material preparation. The novel device allows the deposition of a wide range of powder materials.

Y POWDER MICROFEEDING SYSTEM FOR SOLID FREEFORM FABRICATIONXuesong Lu, Shoufeng Yang, Lifeng Chen and Julian R.G.Evans Department of Materials, Queen Mary, University of London, Mile End Road, London, E1 4NS, UKAbstractSecond generation SFF techniques allow both composition and shape to be downloaded directly from a computer file so that 3D functionally graded materials (FGM) can be assembled. Methods for multi-material deposition are also needed in combinatorial research, colour management and pharmaceutical dosing. In this work, computer-controlled microfeeding systems using ultrasonic vibration of a capillary were built. A wide range of stable flow rate control and switching control were achieved in the acoustic vibration system, and uniform powder doses were obtained in the ultrasonic system. The experimental results show that the nozzle diameter, transmission fluid depth, waveforms, voltage amplitude, frequency and oscillation duration all influence the dose mass.

The experimental results show that the nozzle diameter, transmission fluid depth, waveforms, voltage amplitude, frequency and oscillation duration all influence the dose mass. Among these factors, the nozzle diameter, voltage amplitude and oscillation duration can be used to control the dose mass. Raster printing of patterns with various resolution and dot size are demonstrated.IntroductionIn selective laser sintering (SLS) or three dimensional printing (3DP), layers of powder are laid down, usually with a roller, and the forming areas are either scanned by a high power laser or printed with binders while the non-forming areas remain as loose powder to support subsequent layers [1, 2]. This means the whole object must be made of one material, although in 3DP small amount of doping inks could be printed on the powder bed by changing the ink composition [3] and functional gradient materials (FGM) can be made by depositing a different powder composition in each layer[4].

There is now an interest in incorporating many materials into one component through so-called functional gradients, in which programmed concentration profiles are built in to avoid steep concentration interfaces by concurrent deposition and blending of several different powders. This is becoming possible with direct ink-jet printing [5] and selective laser sintering [6, 7].At the same time, the pharmaceutical industries also need effective techniques for microfeeding of fine powders. Dispensing, dosing and blending of several powders by computer control, allow large potential formulation design spaces. During the early stages of the development of new drugs, the availability of candidates is often limited to a few grams of the newly synthesized powders. For example, the solubility of compounds is one of the screening criteria, and requires automated tools to perform comprehensive early-stage solubility studies.636New methods for dry powder microfeeding are emerging in response.

There are several distinct methods for powder metering and dispensing [8]: pneumatic methods, volumetric dosing, gravimetric dosing, screw and auger dispensers, electrostatic control to meter powders, magnetically mediated flow enhancement for controlled powder discharge of cohesive powders, acoustic or ultrasonic controlled powder dispensers and powder feeding devices based on excitation of a progressive wave in a lossy ultrasonic transmission line. Microfeeding by ultrasonic vibration is one of the promising methods for fine powders.There are a few attempts in the past few years using ultrasonic vibration to dispense dry powder for multi materials solid freeform fabrication. Matsusaka et al. [9 ] first studied the microfeeding of fine powders in a capillary tube vibrated by 20 kHz ultrasound generated by a piezoelectric transducer. Takano and Tomikawa [10] built feeding devices based on excitation of a progressive wave in a lossy ultrasonic transmission line. Li et al.

Takano and Tomikawa [10] built feeding devices based on excitation of a progressive wave in a lossy ultrasonic transmission line. Li et al. [11] studied the mechanism of ultrasonic dispensing. These studies only focus on continuous dispensing with controllable flow rate. Yang and Evans [12,13] studied the factors and mechanism affect the initiation and closure of the flow in vibration controlled dispensing system, which are very important in short dispensing periods. An intermittent dispensing device [14], a dry powder printer, was studied to increase the accuracy of micro-dispensing. Short pulses of ultrasonic vibration were used to dispense the dry powders in “drop” format, analogous to the drop-on-demand ink jet printer. The dispensing mechanism [15], drop uniformity [16], and different design of dispensing nozzles [17] were studied.

The dispensing mechanism [15], drop uniformity [16], and different design of dispensing nozzles [17] were studied. In this paper, we investigate the possibility of pattern printing and drop size control.Experiment detailsThe micro dispensing device comprises a computer, an analogue waveform generator (Model: PCI 6733 DAQmx Card, National Instruments Corporation), an electrical controller (Power amplifier: PB58A, Apex Co., Tucson, Arizona, U.S.A), a piezoelectric ring (Model: SPZT-4 A3544C-W, Size: 35mm(cid:117)15mm(cid:117)5mm, Resonance frequency: 44 kHz, MPI Co., Le Locle, Switzerland), a glass nozzle (capillary tube), a water tank, and a microbalance (2100 mg (cid:114) 0.1 (cid:80)g, Sartorius AG, Goettingen, Germany), as shown in Figure 1. The dispenser was positioned over a three-axis high performance linear motor table (MX80L Miniature Stage, Parker Hannifin Automation, Dorset, UK) capable of high acceleration (39.2 ms-2) and speed (100mms-1).

The table was driven by Labview software (National Instruments, US).The computer and NI 6733 D/A card generate a voltage signal, which can be varied with different waveforms, frequencies and amplitudes. The waveform outputs to the controller, which includes three parts: direct current power supply, power amplification circuit and tuning circuit. The signal voltage was subjected to a gain of 10 when applied to the transducer. The piezoelectric transducer excited by the high frequency signal (>20 kHz) transmits the vibration to the capillary through water. The powder dose mass can be weighed by the microbalance, which transmits the data to the computer via the RS232 serial-port. The inner diameter of the water tank is 40 mm. The feed tubes consisted of glass capillaries drawn down from 10 mm (i.d.) tubes. The upper section acts as the hopper and is used to store the feed powder.

The feed tubes consisted of glass capillaries drawn down from 10 mm (i.d.) tubes. The upper section acts as the hopper and is used to store the feed powder. A piezoelectric ceramic ring was attached637to the bottom of the tank by an adhesive frequently used in ultrasonic cleaning tank construction (9340 GRAY Hysol Epoxi-Patch Structural Adhesive, Dexter Co., Seabrook, USA).1623547RS232 interface1. Computer; 2. D/A card; 3. Electrical controller; 4. Piezoelectric ring; 5. Water tank; 6. Nozzle; 7. Micro balance or 3D tableFigure 1. Experimental arrangement of the microfeeding system.Figure 2. Pattern printing with nozzle mounted on XY table.Results and Discussion1. Effect of powder dispensing parameters on microfeeding The effects of process parameters for different powders have been identified [15].

Effect of powder dispensing parameters on microfeeding The effects of process parameters for different powders have been identified [15]. Dispensing by ultrasonic vibration depends on the powder structures that develop in the capillary tube which can be divided into three types: arching, plugging and blocking. Powders which are cohesive and of low density tend to block the capillary tube as they descend from the hopper, generating many plugs in the capillary. The relatively less cohesive and denser powders cause the arching638phenomenon which allows controlled dispensing. Depending on cohesion, the powder falls as partially compacted columnar rods, clots or cluster or discrete particles, as shown in Figure 3.(a) (b) (c)Figure 3.

Depending on cohesion, the powder falls as partially compacted columnar rods, clots or cluster or discrete particles, as shown in Figure 3.(a) (b) (c)Figure 3. Different drop forms (a) reinforced columnar rod (H13 tool steel) (b) clots or clusters (WC) (c) dispersed discrete particles (glass beads).nozzle diameter, water depth, waveform, voltage amplitude, frequency and oscillation duration all influenced the mean dose mass. Some of these variables need to be fixed and they cannot be used to control dose mass. Feeding should be carried out at the minimum impedance resonance frequency. The water depth affects the resonance frequency of the coupled system and at a given frequency, there is an ideal water depth for inducing the maximum feed. Waveform influences acceleration modes of the piezoelectric element and modifies dispensing; the square wave excitation is preferred for microfeeding.

Waveform influences acceleration modes of the piezoelectric element and modifies dispensing; the square wave excitation is preferred for microfeeding. A range of different nozzle diameters is needed to accommodate fine powders with different densities and particle diameters for microfeeding, there being upper and lower diameter limits for each powder. The minimum dose mass is, in turn, different for different diameter nozzles. Dose mass increases monotonically with excitation amplitude. As the oscillation duration is prolonged, the dose mass increases and for some powders, this relation is almost linear.The experimental results show that microfeeding is best controlled by the voltage amplitude and oscillation duration under square wave actuation. A limited range of nozzles diameters is suited to a given powder. The range of frequency is very narrow for resonance and hence maximum dosing.

A limited range of nozzles diameters is suited to a given powder. The range of frequency is very narrow for resonance and hence maximum dosing. The frequency and water depth are thus fixed for a given dispenser configuration and cannot be used to control the dose mass.2. Drop size on substrate After dispensing, the dry powder dose falls on the substrate and spreads. If the nozzle is close enough to the substrate, the spreading of the dose due to the collision is a minimum. Suppose that the powder piles in a conical shape with radius r and height h. the volume of the cone is (1)where (cid:68) is angle of repose. Then the radius of the dose on substrate is639Ther (cid:32)3m 3 tan (cid:68)(cid:85)(cid:83)(2)where m is the mass of the dose, (cid:85) is apparent density. For WC powder, the angle of repose is 36(cid:113), and apparent density is 8500kgm-3, the 50(cid:80)g dose gives r =0.2mm. The size gives a maximum resolution of 64DPI (Dots per Inch).

For WC powder, the angle of repose is 36(cid:113), and apparent density is 8500kgm-3, the 50(cid:80)g dose gives r =0.2mm. The size gives a maximum resolution of 64DPI (Dots per Inch). The result shows a good fit with the experiment (Figure 4-6). In Figure 4-6, a logo of Queen Mary University of London is printed with 50(cid:80)g doses at different dot pitch and substrate to demonstrate the pattern printing capability. The logo has pixels of 50x50.In Figure 4, the logo was printed with 42 DPI (0.6mm pitch) on plain paper. The dots areclearly discrete, with about 0.1-0.2mm gaps between adjacent dots.1mmFigure 4 Pattern printed with 42DPI (0.6mm pitch) on paper.In Figure 5, the logo was printed with 56 DPI (0.45mm pitch) on alumina slides pre-coated with a thin layer of liquid resin mixed with initiator. The resin has a working time of about 1 hour before polymerization and was used to keep the dots on site after printing.

The resin has a working time of about 1 hour before polymerization and was used to keep the dots on site after printing. Due to the low viscosity and non-uniform thickness of the resin, a few particles drifted from the sites. The resin also causes a slightly bigger dot size due to the drifting of the powders.640Figure 5 Pattern printed with 56DPI (0.45mm pitch) on alumina slide pre-coated with resin.In Figure 6, the logo were printed with 64 DPI (0.4mm pitch) glass slide pre-coated with a thin layer of liquid resin mixed with initiator. The adjacent dots were contact and overlapped with each other giving a dense filling.Figure 6 Powder pattern printed with 64DPI (0.4mm pitch) on glass slides pre-coated with a thin layer of resin.In Figure 7, a 5x5 matrix of dots were printed with increasing dose mass from left to right and from top to bottom.

The matrix was printed with minimum dose mass of about 50(cid:80)g and maximum 1mg, by varying the signal voltage from 1V to 5.8V with increasing step of 0.2V.6415mmFigure 7. Drop size control (from left to right, from the top to bottom, the mass of the dose were varied from about 50(cid:80)g to 1mg, with 25 steps) (f=44.8kHz, t=0.05s).ConclusionsA computer-controlled microfeeding system using ultrasonic vibration of a capillary was built. The effect of process parameters for different powders has been identified. A wide range of stable flow rate control and switching control regime were achieved and uniform powder doses were obtained in the ultrasonic system. Drop size on substrate was estimated through the angle of repose and the apparent density of the powder, and the results concur with the experiments. Raster printing of patterns with various resolutions from 42-64 DPI were demonstrated.

Raster printing of patterns with various resolutions from 42-64 DPI were demonstrated. Dot size can be well controlled through voltage variation.AcknowledgementThe authors gratefully acknowledge the financial support from the Engineering and Physical Sciences Research Council under Grant No. GR/T28812.References1 J.J. Beaman, J.W. Barlow, D.L. Bourell, R.H. Crawford, H.L. Marcus and K.P. McAlea, Solid Freeform Fabrication,A New Direction in Manufacturing, 1-49 and 121-15 Kluwer Academic Publishers, 1997.2 E. Sachs, M. Cima, P. Williams, D. Brancazio and J. Cornie, J. Eng. Ind., 114, 481-488, 1992.3 Yoo J, Cho KM, Bae WS, Cima M, Suresh S: Transformation-toughened ceramic multilayers with compositionalgradients, Journal Of The American Ceramic Society 81 (1), 21-32, 1998.4 J. Yoo, K.M.Cho, W.S. Bae, M. Cima and S. Suresh. J. Am. Ceram. Soc., 81, 21-32, 1998.5 M. Mott, J.R.G. Evans, Zirconia/alumina functionally graded material made by ceramic ink jet printing, Mat.

ying of Colloidal Binder Infiltrated Ceramic Green Parts Produced by Selective Laser Sintering™M. Glazert, N.K. Vailt, and l.W. Barlow+ t. Department of Mechanical Engineering :I: Department of Chemical Engineering The University of Texas at AustinAbstractColloidal ceramic binders have been used to strengthen ceramic green shapes produced by Selective Laser Sintering. This paper focuses on the effectiveness of the colloid infiltration with respect to the physical properties of the colloidal binder. Mass gains, strength gains, and dimensional changes resulting from infiltration were monitored. Controlled drying experiments were conducted to predict the factors influencing drying times for complex shapes. (Key Words: Alumina, Silica Colloid, Drying)IntroductionCeramic preforms can be produced by Selective Laser Sintering of ceramic powders that have been encapsulated with a fugitive polymer binder.

In this process, the polymer binder sinters producing a composite ceramic matrix. However, the polymer binder must be replaced because it does not have sufficient strength nor the thermal resistance necessary for functional ceramic objects. This paper discusses the infiltration of these preforms with ceramic colloids to improve strength and temperature capabilities. Green alumina coupons were infiltrated with colloids of varying solids content and then dried under controlled conditions. Physical properties of the coupons were determined prior to and following colloidal infiltration.MaterialsThe polymer encapsulated ceramic used in this study is the same as that used in a previous study 1. This material consists of a high purity, electronical grade aluminum oxide, provided by Lanxide Corporation, coated with a polymethylmethacrylate polymer having a specified melt flow index of 20g/lOmin as measured by a Kayness Capillary Rheometer at 200°C and 75 psi (ASTM D1238).

Figure 1 shows the mean particle size of the uncoated alumina to be lO.3llm, as measured by a Coulter Multisizer. The particles were observed to be regular in shape using a scanning electron microscope.Samples of Ludox™ colloidal silica, grade TM, were provided by Dupont Corporation. This colloid is 50% wt. silica and consists of particles having a mean diameter of 22nm. A sample of Dupanol ME, a sodium dodecyl sulfate based emulsifier, was also provided by Dupont Corporation.MethodsSLS Production of Test CouponsTest specimens were produced using an SLS Model 125 workstation equipped with a 25 watt C02 laser. Fabrication of parts was conducted in an inert nitrogen environment using the333parameters listed in Table 1. A number of rectangular coupons I l xl l xO.25" (Group 1), I l xl l xO.5" (Group 2), and 1.25"x1.25"xO.25" (Group 3) were produced for drying studies. Additional coupons 1"x3"xO.25" were produced for strength measurements.

Additional coupons 1"x3"xO.25" were produced for strength measurements. Densities of these coupons were determined by mass and dimensional measurements.100 .--------------~-~80..........................................................•....•.......................................Dp:::: 10.3 ± 5.6 J.III120W1 Particle Diameter, 4tm)~nFigure1. Particle size distribution of Lanxide alumina.Table I. SLS panuneters for test specimens used in this study.104250-11065Post-processing of preformsEach coupon size was infiltrated with colloids of varying silica content. The stock colloid was diluted with deionized water to produce samples with 10%,20%,30%, and 40% wt. silica. Emulsifier was added to each colloid sample to aid infiltration in an amount of 1.0% wt. based on total water content.Infiltration was accomplished by placing a coupon in a pool of colloid and allowing it to draw up the colloid.

based on total water content.Infiltration was accomplished by placing a coupon in a pool of colloid and allowing it to draw up the colloid. When the colloid reached the top of the coupon additional colloid was dripped on to the top surface. The coupon was then transferred to a paper towel and excess colloid was drawn off. When no further colloid could be drawn off, colloid was dripped on the top of the coupon. Infiltration was considered complete when a wet spot appeared on the paper towel.Following infiltration the coupons were weighed, measured, and transferred to a drying chamber. Figure 2 shows the constant humidity drying chamber used in this study. Temperature and relative humidity in the chamber were monitored with a thermohygrometer. Temperature was maintained at the room temperature of 24°C. Relative humidity was maintained at 64±2% with a sodium dichromate salt solution2.

Temperature was maintained at the room temperature of 24°C. Relative humidity was maintained at 64±2% with a sodium dichromate salt solution2. An air flow rate of 45 ft/s, measured by a Alnor Velometer, was provided by an electric fan controlled by a rheostat. Laminar air flow was created by forcing the air through a group of narrow, adjacent tubes.In the chamber, the coupons were placed on a platform suspended from a cantilever balance. The balance was connected to an LVDT which sent a signal to a data acquisition system. The initial and final masses of the sample provided the calibration of the LVDT. The data was used to observe the changing moisture content of the part as a function of time. After drying, the bars were removed from the chamber, weighed, measured, baked at 200°C for 1 hr.

The data was used to observe the changing moisture content of the part as a function of time. After drying, the bars were removed from the chamber, weighed, measured, baked at 200°C for 1 hr. to remove residual water and then weighed again.334Results and DiscussionPrior to conducting drying tests, a set of experiments was performed to verify the effectiveness of the colloidal infiltration described above. This was done by introducing a dye into the colloid and observing cross~sections of wet coupons. Discoloration of cross~sections was observed to be homogeneous. Furthermore, x~ray mapping of cross~sectionsbyscanning electron microscopy showed silica to be present at the center~line and a slight concentration gradient to exist. These two methods proved the infiltration method to be effective. Figure 2 shows SEM micrographs comparing pure alumina powder to silica infiltrated alumina.

These two methods proved the infiltration method to be effective. Figure 2 shows SEM micrographs comparing pure alumina powder to silica infiltrated alumina. The figures clearly show an accumulation of material in the infiltrated sample.Figure 4 shows resultant densities of square coupons. Prior to colloid infiltration, all coupons had an initial density of approximately I.SSg/cm3. As would be expected coupon density increases with increasing silica content in a nearly linear fashion. From mass balance considerations the observed increases in density are a result of silica infiltration.

At the critical moisture content, Olc, the rate falls as the drying front recedes into the part. Analysis of these curves for the coupons studied here indicate the drying rate to be influenced by both coupon thickness and the solids content of the infiltrant. Using the analysis of Perry and Green4, it was determined that the drying process is controlled by diffusion and capillary moisture transport. A rigorous analysis of the data is in progress.Cracking of coupons was observed to increase as the thickness of the part increased. Cracking also increased as the solids content of the infiltrant increased. In a separate set of experiments, coupons, that were infiltrated with waterthen dried under similar conditions, did not develop cracks. Cracking appears to be related to stresses induced by drying of the colloid and to the initial strength of the green coupon. Scherer5 has shown that ceramic colloids tend to shrink and crack during controlled drying.

Scherer5 has shown that ceramic colloids tend to shrink and crack during controlled drying. Therefore, as the drying front proceeds in to the coupon, stress gradients that ex.ceed the green strength of the coupon may develop Thicker coupons would necessarily experience larger stress gradients and be more likely to crack.To examine the influence of initial green strength a set of coupons with similar dimensions but with different strengths were processed. Coupons having strengths of 45psi, 100psi, and 160psi were infiltrated with 30 wt. % colloid and dried under similar conditions. The low strength bar cracked severely, the high strength bar did not crack, and the moderate strength bar developed hairline cracks.337Conclusions/Further WorkInfiltration with silica colloids is an effective method for increasing strength and density of polymer. bound.

bound. prototypesprodl1cedby.SelectiveLayerSintering.$trength.gainsof the coupons were nearly independent of the solids content of the infiltrant. However, high solids infiltrants increased cracking, especially in parts with low inherent green strength.Further study is required to optimize the infiltration and drying processes. Additional strength measurements need to conducted on coupons from which the polymer binder has been removed. It may be necessary to subject the coupons to an additional colloid infiltration following binder removal to further increase density and strength. The strength of fired bars need to be studied, also.A detailed analysis of stresses during the drying period, which includes the contributions of Infiltration with other ceramics couldcapillary and diffusion induced drying, is in progress.

Both paths utilize a circular contour with serpentine hatch fill; however, successive layer patterns are varied and the effects compared between 90° and 120° angular pattern shifts. During fabrication, the part’s thermal history and melt pool temperature are recorded via an in-chamber infrared (IR) camera and a dual-wavelength (DW) pyrometer, respectively. These tools are used for non-destructive thermographic inspection (NTI) of the part to ensure target quality and/or microstructure. A unique calibration method for the IR camera utilizing the DW pyrometer data is presented and a calibration correction factor was utilized for high temperature ranges. The melt pool was found to be 40-50% superheated reaching temperatures up to 2500 ºC at times. Temperature characteristics of two different layers were compared for different hatching patters, and the results show that for a given point in time, maximum temperatures can vary based on laser raster.

Temperature gradients varied and peaked at about 1000 ºC/mm along the diameter of the small rods. This can lead to anisotropy in microstructural and mechanical properties allowing for unique property growth per build path. Cooling rates within the melt pool appear to increase as maximum melt pool temperature increases, for instance, from 16,000 ºC/s – 41,000 ºC/s.IntroductionLaser Engineered Net Shaping (LENSTM) is a Direct Laser Deposition (DLD) process first developed by Sandia National Laboratories [1]–[4]. This laser-based additive manufacturing (LBAM) technology has the potential to reduce production costs by lowering the amount of time and material required for fabrication. Direct Laser Deposition utilizes a slice file created from a 3-D Computer Aided Design (CAD) solid model to deposit a geometry one layer at a time. Each layer is deposited in tracks by feeding metal powder into a laser-formed melt pool onto a substrate.

Each layer is deposited in tracks by feeding metal powder into a laser-formed melt pool onto a substrate. Before the benefits of DLD can be fully realized, a firm knowledge base must be developed to characterize the effects of DLD process parameters - such as laser power, travel speed, and powder feed rate - on a part’s thermal history, which in turn determines its final microstructure and mechanical properties.The length and maximum temperature of the melt pool increases with laser power and slower travel speeds. As the melt pool moves, its tail rapidly solidifies and experiences the highest cooling rates [5] and these rates increase with travel velocity [6]. During DLD, delivered laser energy is transferred by conduction away from the melt pool through previous layers and the substrate while the remainder contributes towards melting and vaporization or is transferred away from the surface by convection and radiation [7].

Maximum DLD Ti-6Al-4V melt pool temperatures can be up to 40% above their liquid temperature (i.e. superheat). Furthermore, local melt pool temperature gradients and peak cooling rates for parts undergoing DLD have been experimentally and numerically assessed at 100 °C/mm – 1000 °C/mm [3], [8]–[11] and 12,000 °C/s – 25,000 °C/s [12], [13] .For the uncontrolled DLD process, in which time-invariant process parameters are employed, the final part microstructure is greatly prone to anisotropy. This is due to many thermal-fluidic phenomena, including: localized heating from the laser, non-uniform thermal gradients, solidification and time-dependent cooling rates. Better understanding of the sensitivity of such thermal phenomena to the uncontrolled DLD process parameters can allow for their easier manipulation for minimizing anisotropy.

Better understanding of the sensitivity of such thermal phenomena to the uncontrolled DLD process parameters can allow for their easier manipulation for minimizing anisotropy. One way of controlling heat transfer, and thus final product quality, in both LENS and other DLD or cladding processes, is to use nondestructive thermographic inspection (NTI) for generating data to be used in feedback control [4], [14]–[16]. For example, one can relay melt pool temperatures to a laser controller for microstructural uniformity throughout a part by compensating laser delivery based on local heat transfer [16]. Common tools used for LBAM monitoring include infrared (IR) cameras and dual- wavelength (DW) pyrometers [2], [3], [17]–[19].In order to effectively utilize an IR camera for monitoring absolute temperature of a part during DLD (or laser cladding), the spectral/directional, temperature-dependent surface emissivity must be known.

Due to radiation noise in a typical DLD chamber, and the transient nature of surface emission, determining the surface emissivity can be very challenging if not impossible. Thus, in-lieu of determining the absolute temperature distribution of a part, one can opt to utilize relative temperature gradients provided by IR images. However, Griffith et al. suggested that, if a valid expression for effective wavelength (as a function of temperature) is known, then one can employ Wein’s Law to determine absolute temperature [8].There has been some recent work focused on calibrating IR cameras for use in LBAM processes. In 2005, Yang et al. showed that the emissivity of Ti-6Al-4V remained relatively constant (at ~0.25) between 132 °C to 800 °C after which point it varied up to 0.98 at 1029 °C due to severe oxidization [20]. Doubenskaia et al. experimentally determined the emissivity of Ti-6Al-4V during a Nd:YAG laser cladding process purged with argon [21].

Doubenskaia et al. experimentally determined the emissivity of Ti-6Al-4V during a Nd:YAG laser cladding process purged with argon [21]. By using a MICRON M390 to find brightness temperature with the IR camera, an emissivity of 0.201 was determined in the phase transition temperature of the metal. Actual temperature was then found by using the brightness temperature at liquid-solid transition, the determined emissivity, and the grey body assumption [21]. Rodriquez et al. determined the emissivity of Ti-6Al-4V during a Selective Laser Melting (SLM) process to be 0.26 between 650 °C and 750 °C (similar to [20] and [21]) and 0.5 in powder form [22]. These values along with reflected temperature and optics transmission were used to calculate actual surface temperature which was validated by usingthermocouples during the build.

These values along with reflected temperature and optics transmission were used to calculate actual surface temperature which was validated by usingthermocouples during the build. Actual temperatures varied from the recorded IR temperatures by a near-constant value.Pyrometers are an attractive means for non-contact temperature measurements in DLD processes as they are able to measure high temperatures (>1000 °C) [16]–[19], [23], [24]. They are, therefore, ideal for measuring melt pool temperatures that can exceed the liquidus temperature of, say, Ti-6Al-4V – at 1636 °C. A DW pyrometer may be selected to reduce the error caused by sensitivity to emissivity [13]. Utilization of two wavelengths allows for more accurate surface temperature measurements by using the ratio of the relative radiation intensity at both wavelengths [12], [13].

Utilization of two wavelengths allows for more accurate surface temperature measurements by using the ratio of the relative radiation intensity at both wavelengths [12], [13]. Many have used thermal monitoring for characterizing maximum temperatures and cooling rates of a bulk part and melt pool during DLD, and relating them to the microstructure and mechanical properties of parts [2]–[4], [8], [15], [17], [19], [25]. For instance, Kobryn et al. mapped out solidification rates using thermography while experimenting with various lasers in an additive cladding process. Solidification rates were then used to justify the microstructure seen in the final parts [25]. Nassar et al.

Solidification rates were then used to justify the microstructure seen in the final parts [25]. Nassar et al. related thermal history to the growth of microstructure and hardness along the height-wise direction of a build [16].This current work focuses on utilizing LENS for fabricating Ti-6Al-4V cylindrical specimens while monitoring their bulk and melt pool temperature responses, via in-chamber IR camera and DW pyrometer, respectively. A demonstration of dual thermography of cylindrical specimens during LENS and a unique approach for calibrating the high-end temperature of an IR camera using online pyrometer data is provided. Thermographic images of the melt pool and bulk part are provided. The effect of scan strategy (i.e.

Thermographic images of the melt pool and bulk part are provided. The effect of scan strategy (i.e. build pattern) on the temperature and melt pool response is discussed, in addition to typical heating and cooling rates and temperature gradients in vicinity of the heat affected zone (HAZ).Experimental Setup and ProcedureExperiments were conducted to observe build pattern effects on part/pool thermal signatures at various layers, as well as to determine a correction factor (CF) for the IR camera for DLD of Ti-6Al-4V. An Optomec LENSTM 750 DLD system (chamber shown in Figure 1) with a 1 kW Nd:YAG laser was used for part fabrication. Titanium alloy (Ti-6Al-4V) powder (Phelly Materials Inc.), manufactured via a Plasma Rotating Electrode Process (PREP) with -100/+325 mesh size was used.

Titanium alloy (Ti-6Al-4V) powder (Phelly Materials Inc.), manufactured via a Plasma Rotating Electrode Process (PREP) with -100/+325 mesh size was used. The powder was deposited onto a 152.4 x 152.4 x 3.175 mm3 Ti-6Al-4V substrate placed atop a 6.4 mm thick copper spacer-plate that was attached to a moving build stage.The following parameters were utilized (values approximate): 350 W laser power, 16.9 mm/s scan speed, and 0.13 g/s powder feed rate. The DLD chamber was sealed and purged with argon during all builds. Cylindrical specimens, Ø = 6.6 mm and 78.2 mm tall, were fabricated using a 5.08 mm layer thickness and hatch spacing (each specimen was a total of 154 layers). Two different build strategies, characterized by unique hatching (contour fill) patterns (Pattern A and B) were examined in this study.

Two different build strategies, characterized by unique hatching (contour fill) patterns (Pattern A and B) were examined in this study. For both patterns, the outer circumference of the rod was deposited first in a counter-clockwise direction, followed by alternating tracks to fill in the circle as shown in Figure 2. Pattern A designates a 0°, 90°, 180°, and 270° repeating sequence of the aforementioned track deposits while Pattern B designates a 0°, 120°, and 240° repeatingsequence. It is of interest to note that Ensz et al. demonstrated that a build strategy consisting of 105° track rotations between layers can be optimal for fabrication of cylindrical parts via LENS [26]. However, this current study is more concerned with the thermal response measurement, and not the optimization, of hatch patterns.Deposition HeadIR CameraCylindrical SpecimenTi-6Al-4V SubstrateBuild StageCopper SpacerFigure 1.

Photograph of Optomec LENS 750 chamber with deposition head, substrate, spacer and custom-installed IR camera.(a)Contour is deposited first(b)Figure 2. Contour deposit and hatch raster pattern and order for LENS: (a) Pattern A: 90° rotations for each layer and (b) Pattern B: 120° rotations for each layer.As shown in Figure 1, an IR camera (Stratonics) was attached to the build stage within the LENS chamber to allow for monitoring of the exposed part surface and HAZ during the build. Using this setup, the deposition head appeared to be moving during visualization as the tracks were deposited. The IR camera was positioned approximately 101.6 mm above the substrate and pointed 10º (from vertical) towards the build, as shown in Figure 3, which also shows the focal plane and a few layer numbers. During experiments, the IR camera was set to capture images at 14 Hz.

During experiments, the IR camera was set to capture images at 14 Hz. In addition to the IR camera, a DW pyrometer (Stratonics), located atop and outside the LENS chamber, was utilized for monitoring melt pool temperature (and indirectly, melt pool geometry and size). As shown in Figure 3, the DW pyrometer was mounted in such a way that its line-of-sight was directed down the deposition head column to the melt pool. The DW pyrometer possessed a 16 micron/pixel resolution and a 4 mm field of view. Using Stratonics ThermaViz software, the DW pyrometer was set to capture images while in the ‘free-run’ setting, resulting in an acquisition frequency between 4 - 7 Hz.

Using Stratonics ThermaViz software, the DW pyrometer was set to capture images while in the ‘free-run’ setting, resulting in an acquisition frequency between 4 - 7 Hz. For calibration and comparison, melt pool and bulk temperature data were extracted from a ‘measurement line’ along the left edge (relative to camera point of view) of the rod as shown in Figure 3.Pyrometer Line-of-SightIR Focal PlaneLayer 15410°Measurement LineIR CameraLayer 15Cylinder BuildLayer 1SubstrateFigure 3. Side view of the IR camera focal plane which is oriented ~10° from vertical, including: arbitrarily-selected layer numbers, DW pyrometer line-of-sight, fictitious measurement line, completely-built cylinder and the substrate.Infrared Camera CalibrationThe as-received DW pyrometer was calibrated using a high-temperature tungsten filament at known temperature.

Since DW pyrometers provide for relatively accurate temperature measurements at each point along a layer during metallic deposition [20], the in- phase DW pyrometer and IR data, collected along the measurement line, were utilized for high- temperature calibration of the IR camera. The in-phase, maximum IR measurements and average melt pool temperatures at the measurement line were employed; corresponding to the deposition of the circular contour for each seventh layer. The melt pool temperature was quantified using the mean of all values greater than the liquidus temperature of Ti-6Al-4V (1636°C [27]) along fictitious lines parallel and perpendicular to the melt pool’s maximum temperature zone. Absolute maximum temperature of the melt pool was also recorded. As shown in Figure 3, the cylindrical rods were placed on the center of the substrate to ensure their relative position to the IR camera focal plane was known with reasonable accuracy.

As shown in Figure 3, the cylindrical rods were placed on the center of the substrate to ensure their relative position to the IR camera focal plane was known with reasonable accuracy. Note that, for the current setup, the IR temperature error increases with layer number due to the focal plane diverging from the measurement line.Figure 4 provides the in-phase maximum, raw IR measurements and separate melt pool temperatures (from the DW pyrometer) as recorded upon intersection of the melt pool (during the initial layer contour deposit) with the measurement line while employing either Pattern A or B.

For the calibration, the corresponding, maximum melt pool IR signature was assumed equivalent to the average temperature of the melt pool from the pyrometer.Pattern A Melt Pool Maximum Pattern A Average of Melt Pool Pattern A Max IR SignalPattern B Melt Pool Maximum Pattern B Average of Melt Pool Pattern B Max IR Signal30002500)C ° ( e r u t a r e p m e T200015001000500018 15 22 29 36 43 50 57 64 71 78 85 92 99 106 113 120 127 134 141 148 155Layer NumberFigure 4. Comparison between in-phase IR signal of melt pool at measurement line and either maximum or average melt pool temperature (measured via pyrometer) as measured during Pattern A and B deposits.From Figure 4, it is clear that both build patterns have similar average melt pool temperatures and IR values throughout the build; however, the melt pool’s maximum temperature seems to vary more since the starting point of the contour deposit can vary between both build patterns.

If a measurement was taken at the beginning or end of a deposited contour where the melt pool velocity approaches zero, the melt pool would be expected to have a higher temperature [6]. The maximum melt pool temperature is found to vary between ~ 2,100 ºC and 2,500 ºC during contour deposition and be almost independent of deposition pattern. The average melt pool temperature was found to vary between 1,900 ºC and 2,000 ºC during contour deposition. These pyrometer measurements indicate that the Ti-6Al-4V melt pool can be 20 - 50% superheated above its liquidus temperature. Note that both the melt pool temperature and IR signal were lower during the deposition of initial layers.Each raw, uncorrected IR data point was corrected to absolute temperature via summation of a correction factor [22], and this method is herein utilized by use of pyrometer data shown in Figure 4.

A correction factor (CF) of CF ≈ 886 was found by taking the average difference between each raw IR data point and its corresponding pyrometer temperature measurement for Pattern A. Since the emissivity of molten Ti-6Al-4V varies only slightly in an inert environment [20], [22], the CF was assumed constant regardless of temperature. This appears to be an accurate assumption, since, as shown by Figure 5, the difference between the maximum IR signal and average melt pool temperature is approximately constant for both deposition patterns. The percent relative error between pyrometer/melt-pool temperature and post-calibrated IR temperature is provided in Fig. 5.

The percent relative error between pyrometer/melt-pool temperature and post-calibrated IR temperature is provided in Fig. 5. It can be seen that the average, relative error is approximately 2%, with maximum error occurring during deposition of Layer 1, corresponding to 5.9% for Pattern A and 14.6% for Pattern B.Pattern A Pattern B Correction Factor (CF) Error from Applied CF Pattern A Error from Applied CF Pattern B1400)C ° ( e r u t a r e p m e T120010008001412108n i e c n e r e f f iD6004002006420018 15 22 29 36 43 50 57 64 71 78 85 92 99 106 113 120 127 134 141 148 155Layer NumberFigure 5. Difference in measurement and percent relative error between calibrated, maximum IR temperature and average melt pool temperature for Pattern A and B.Results & DiscussionThe melt pool and part temperature response during contour deposition for Layer 85 and Layer 148 are shown in Figure 6 and Figure 7, respectively.

Note that the white lines in Figure 6 and Figure 7 represent the line along which a measurement was taken. All pyrometer images (melt pool) are shown on a scale from 1400 °C – 2400 °C while the IR images are shown on the same scale only for the enclosed calibrated regions.)%(r o r r ETemperature (°C)(a)Calibrated Region(b)Calibrated RegionFigure 6. Thermographic images (melt pool captured by pyrometer on left and IR temperature distribution on right) collected simultaneously during contour deposit of Layer 85 for (a) Pattern A and (b) Pattern B. For the bulk temperature distribution, the scale is valid for the dashed box region only.Calibrated RegionCalibrated RegionTemperature (°C)(a)(b)Figure 7. Thermographic images (melt pool captured by pyrometer on left and IR temperature distribution on right) collected simultaneously during contour deposit of Layer 148 for (a) Pattern A and (b) Pattern B.

For the bulk temperature distribution, the scale is valid for the dashed box region.For Pattern A, the maximum melt pool temperature was slightly higher than for Pattern B in Layer 85 – a difference of about 52 °C, but the average temperature for both was about 1940 °C. The general shape of the melt pool temperature contour is also similar for both patternsshown in Figure 6. Both build patterns were laying their tracks at 0° rotation and were thus identical; therefore, one would expect such similar results from the thermal data. A 121 °C and 73 °C temperature difference exists between the maximum melt pool and average melt pool temperatures respectively at Layer 148 as shown in Figure 7 – with Pattern B being the greater. The diameter of the melt pool is also smaller here in Pattern B and suggests that the laser has just begun its contour movement; whereas, Pattern A seems to have been half-way through its deposit.

The diameter of the melt pool is also smaller here in Pattern B and suggests that the laser has just begun its contour movement; whereas, Pattern A seems to have been half-way through its deposit. Since the laser’s velocity at the beginning of the contour is zero, this creates a build-up of heat and a more concentrated melt pool. The melt pool begins to cool and lengthen as the velocity is increased as Figure 7 (a) suggests. At this point in time, Pattern B is laying a 0° rotation track while Pattern A is laying a 270° rotation track; therefore, one could expect a variation in temperature measurements as evidenced in the data.

Though these maximum and average temperatures are well above any phase transformation range, their difference in magnitude still agrees that variations in microstructural and mechanical properties can exist within a part under certain conditions due to temperature.Cooling rates were found by taking the spatial derivative along the direction of melt pool travel, i.e. subtracting neighboring temperatures and dividing by the distance between them. The following equation may be applied to find cooling rates within the melt pool [13]:𝐶𝑟(𝑡) =∆𝑇𝑚𝑝 ∆𝑥𝑎𝑣𝑔𝑠where 𝐶𝑟(𝑡) is the cooling rate as a function of layer; ∆𝑇𝑚𝑝 is the change in temperature between averaged points in the melt pool that are an average distance of ∆𝑥𝑎𝑣𝑔 apart; and s is the laser scan rate equal to 16.9 mm/s. The locations of greatest slope, i.e. maximum cooling rate, were found at the tailing end of the melt pool as a preliminary step. Figure 8 gives an example of such location.

The locations of greatest slope, i.e. maximum cooling rate, were found at the tailing end of the melt pool as a preliminary step. Figure 8 gives an example of such location. Next, three temperature points taken from the DW pyrometer data for a single melt pool image were averaged on both sides of the location of maximum temperature change within the melt pool and then subtracted from one another to give a ∆𝑇𝑚𝑝. An optical calibration determined the distance between pixels to be about 0.014 mm making ∆𝑥𝑎𝑣𝑔= 0.042 mm when the locations corresponding to the temperatures on either side of the maximum change location were averaged and subtracted.2400Slope of maximum cooling rate)C ° ( e r u t a r e p m e T22002000180016001400Travel directionMelt pool leading edge12000.001.00 Parallel Melt Pool Location (mm)0.501.502.00Figure 8. Maximum cooling rate (slope) for Layer 85 of Pattern A.

Maximum cooling rate (slope) for Layer 85 of Pattern A. Data averaged on either side of location of maximum slope.Values for the cooling rates are shown in Table 1 as well as the pattern rotation deposited at the current layer number. The data shows that the first layer has the lowest cooling rate. Profiles of the melt pool at this first stage are still developing; therefore, the calculated slopes were less in this layer than the other observed layers. Both patterns experience an increase in cooling rate in the middle layers. At these points, the melt pool has had time to develop – creating larger cooling rates within the melt pool temperature profile. Pattern B’s cooling rate is still greater in Layer 148 compared to Layer 85 – its maximum temperature is increasing (see Figure 4), and the temperature slopes become steeper. However, a comparison between Pattern A’s Layer 148 and Layer 85 shows a decrease in maximum temperature.

However, a comparison between Pattern A’s Layer 148 and Layer 85 shows a decrease in maximum temperature. Along with the temperature decrease, the pattern is also different. As previously mentioned, the data suggested that Pattern B was near the beginning of its contour deposition which created elevated temperature over Pattern A. With similar average melt pool temperatures, an increase in maximum melt pool temperatures corresponds with an increase in cooling rate. Others have seen cooling rates of 25,000 °C/s [13] and 12,000 °C/s [12]. These cooling rates are comparable to tabulated in Table 1; however, the increased scan speed as compared to [13] has caused these results to exceed 25,000 °C/s.These melt pool cooling rates should not be confused with bulk cooling or solidification rates that will exist in a fabricated part. For instance, bulk cooling could be expected to be greater in the first few layers as the substrate acts as a heat sink.Table 1.

For instance, bulk cooling could be expected to be greater in the first few layers as the substrate acts as a heat sink.Table 1. Estimated cooling rates associated with three different layers for each pattern.Pattern APattern BLayer Number1 85 148Cooling Rate (°C/s) 15,661 41,675 24,156Pattern Rotation (°) 0 0 270Cooling Rate (°C/s) 16,060 33,844 41,144Pattern Rotation (°) 0 0 0The temperature gradient along six layers measured by the IR camera is shown in Figure 9. These measurements were taken when the melt pool was located relative to the measurement line in the IR image of Figure 3. Temperature distributions per layer were very similar between the two patterns – the maximum signal reached throughout the build was close to 1986 °C. Since temperature is shown as estimated actual temperature, only data with an IR signal above 800 °C (the minimum range of this study) has been considered and calibrated.

Since temperature is shown as estimated actual temperature, only data with an IR signal above 800 °C (the minimum range of this study) has been considered and calibrated. Location along the diameter of the rod was found by first noting the pixel at which IR signal began to increase. Then, the last point at which the temperature decreased was considered to be at the outer diameter of the part (6.6 mm). Distance between pixels was then found to be about 0.367 mm. Discrepancies between each patterns’ contours of Layers 113 and 154 respectively may be attributed to the IR camera being slightly out-of-focus for these layers.Spatial temperature distribution data from the IR camera can be derived by subtracting neighboring temperature values and dividing by the distance between pixels – about 0.367 mm – to create the plots shown in Figure 10.

Temperature profile measured by the IR camera along the diameter of the rod at the melt pool for different layers of (a) Pattern A and (b) Pattern B.(a)(b)Layer 43Layer 85Layer 113Layer 43Layer 85Layer 113Layer 148Layer 154Layer 148Layer 154)m m C ° (/t n e i d a r G1000800600400200)m m C ° (/t n e i d a r G1000800600400200l a i d a R d e t a mi t s E0200400600800l a i d a R d e t a mi t s E020040060080002460246Location on Rod Diameter (mm)Location on Rod Diameter (mm)Figure 10. Plot of estimated radial gradients against location on rod diameter for (a) Pattern A and (b) Pattern B.ConclusionsUsing dual-thermography for non-destructive thermal inspection (NTI) of fabricated parts during Directed Energy Deposition (DED) processes such as Laser Engineered Net Shaping(LENS) will be important for assessing part quality and microstructure in real-time.

In a pyrometer/IR-camera setup, the melt pool temperature can be utilized for calibrating the IR camera which can be located in the LENS chamber. In this case, a near-constant calibration offset is feasible for a given temperature range. The current results indicate that the IR camera high-temperature can achieve accuracies on-the-order of 2-5 %.Build-path temperature effects were not visibly discernable when varying the scan patterns for the cylindrical layers. However, upon analyzing thermographic data, temperature gradient magnitudes and cooling rates were found to vary along the radius of the melt pool and cylinder and this can influence final microstructure. For a given point in time and space, top- level layers were shown to have temperature characteristics more dependent on hatch pattern. Maximum melt pool temperature was observed to reach as high as 2500 ºC and to be dependent on its contour track’s starting point.

Maximum melt pool temperature was observed to reach as high as 2500 ºC and to be dependent on its contour track’s starting point. The increase in melt pool temperature appeared to translate to an increase in cooling rate from about 16,000 ºC/s - 41,000 ºC/s for both investigated hatching patterns. Radial temperature gradients decreased with increasing layers; ranging between 500 ºC/mm – 1000 ºC/mm near the tail-end of the melt pool. The hatch pattern consisting of less layer revolutions appeared to have higher radial temperatures gradient, perhaps due to a mismatch of the patterns’ deposition tracks.Future work will aim to relate the observed thermal phenomena to determined microstructures using cooling rates based on the tail of the melt pool. A larger temperature range will be calibrated so that the IR camera may be more readily used for determining post solidification evolution.

al-Material Minimum Weight Structures Fabrication Using Ultrasonic ConsolidationObielodan J.O1 and Stucker B.E2*1Mechanical and Aerospace Engineering Department, Utah State University, Logan, UT 843222Industrial Engineering Department, University of Louisville, Louisville, KY 40292AbstractThe multi-material capability of additive manufacturing (AM) processes has createdopportunities for structural designs that would otherwise be impossible. This work involves thedevelopment of a methodology for fabricating dual-material minimum-weight structures usingultrasonic consolidation (UC). Sample structures were designed, fabricated and tested for loadcarrying capabilities. Analyses of results show that dual-material minimum weight structuresmade of Al3003/MetPreg® and Al3003/Ti composite material members can withstandsignificantly higher strain energy densities up to the point of failure than similar structures madeof Al 3003 alone.

This is an indication that UC can be effectively used to fabricate multi-materialstructures for real life applications.1IntroductionResearch on the fabrication of multi-material structures using different additivemanufacturing (AM) processes has accelerated in recent years. AM processes have the potentialfor flexible variation of materials and microstructures, both in continuous and discrete fashion inaddition to their capabilities for complex geometry structures. New uses of some advancedmaterials are being discovered because of the ability to combine them with other materials inAM fabricated structures. Also, some of the materials that would otherwise be difficult tocombine in conventional processes are being processed (Cohen et al., 2006; Malone et al., 2004;Liu and DuPont 2003; Griffith et al., 1997; Arcaute et al., 2009; Janaki Ram et al., 2007;Obielodan et al., 2010B).

This ability to deposit function specific materials where they areneeded in a structure further revolutionizes engineering structure design and material usage. Oneof the driving forces is the economic use of costly advanced materials that are prescribed to bedeposited just where they are functionally required in a structure. The application of thesecapabilities is diverse, ranging from medical to aerospace, automobile, nuclear and others.Potential applications of multi-material structures fabricated using AM processes havebeen demonstrated. Arcaute et al., 2009, used stereolithography (SL) for the fabrication of multi-material scaffolds with spatially controlled characteristics for tissue engineering applications.Also, Wicker et al., 2004, fabricated complex multi-material hydrogel constructs for nerveregeneration and guided angiogenesis applications. Meso and macro scale multi-materialstructures have also been fabricated using SL (Jae-Won Choi et al.

Meso and macro scale multi-materialstructures have also been fabricated using SL (Jae-Won Choi et al. 2009 and Inamdar et al.2006).Objet Geometries Limited commercialized the 3D printing of dissimilar material end useproducts using polymer materials (www.growit3d.com/services/multi-material-polyjet). Objet’sConnexTM machines jet multiple materials simultaneously to fabricate multi-material structures.Different laser powder metal deposition processes have been used to deposit multi-material structures. Examples include gradient structures (Griffith et al., 1997, Liu and DuPont,2003), surface cladding with corrosion and wear resistant materials for machinery (Foroozmehret al., 2009) and medical implants applications (Janaki Ram and Stucker, 2008).

Both 3Dprinting and laser powder deposition processes have capabilities for continuous materialvariation as well as discrete material domains in fabricated multi-material structures.Ultrasonic consolidation has been demonstrated to have the capabilities for multi-materialstructures fabrication. This capability was demonstrated by Janaki Ram et al., 2007B, in theirwork in which copper, brass, nickel, inconel 600, AISI 347 stainless steel, stainless steel AISI304 wire mesh, MetPreg®, and aluminum alloy 2024 were each welded to aluminum alloy 3003-H18. Domack and Baughman, 2005, investigated the capability of UC to fabricate gradedtitanium and nickel alloy multi-material structures. Additionally, ultrasonic welding has beensuccessfully used to weld metals to a polymer matrix composite (Kruger et al., 2004).

Additionally, ultrasonic welding has beensuccessfully used to weld metals to a polymer matrix composite (Kruger et al., 2004). Obielodanet al., 2010B, further demonstrated UC multi-material capabilities by welding differentcombinations of molybdenum, tantalum, titanium, copper, silver, nickel, MetPreg®, aluminumalloys 1100, 3003, 6061 and boron powder. Also, the shear strengths of titanium/aluminumultrasonically bonded foils were characterized by Obielodan et al., 2009.Ultrasonic consolidation, described more fully elsewhere (White, 2003 and Obielodan etal., 2010A), is a low temperature process that combines ultrasonic welding and additivemanufacturing technology. Multi-material structures fabricated using UC characteristically havediscrete material domains as opposed to the continuous material variation that is obtainable withlaser powder deposition processes and 3D printing.

The process has the potential for fabricatingstructures for applications in systems subjected to mechanical loading. In this study, amethodology for fabricating multi-material structures using UC was developed. UC structureswith a single material are relatively easy to fabricate when compared to multi-materialfabrications, as foils can be automatically fed.Multi-material minimum weight Michell structures (Dewhurst, 2001; Dewhurst, 2005;Selyugin, 2004 and Michell, 1904) represent one of the categories of structures that can begeometrically and materially complex to fabricate using conventional processes. They are madeof multiple, thin members that are preferably made from light weight materials with high specificstrength and stiffness. Such structures are readily applicable to aerospace and automotiveindustries, where there is continuous emphasis on higher strength and lower weight structures forimproved fuel efficiency and performance.

Figure 1 shows an example of a complex minimumweight structure with members that could be made of different materials based upon Michelltheory (Michell, 1904). In the illustrated structure, if the structure is pinned at points A and Band a load is placed at C, parallel to a line between A&B, as shown with the arrow, the outermember labeled D will be in pure compression, as well as all the inner members that join Dtangentially. Those inner network members that are perpendicular to D, and the memberbetween A and C will be in pure tension. In order to optimize a structure to its fullest extent, themembers in tension can be made of materials different from those in compression.

In order to optimize a structure to its fullest extent, themembers in tension can be made of materials different from those in compression. In this case,the intersection between the tensile and compressive members and the design and strength ofthese joints is of critical importance for the structure’s reliability and performance.ABDCFigure 1: A minimum weight structure design (Dewhurst, 2001)Simplified minimum weight structures based on maximum strength and maximumstiffness criteria (Dewhurst, 2005) were designed and fabricated. Figure 2 shows a free bodydiagram of the structure with oa, ob and oc as compression members and ab and ac as tensionmembers when subjected to compressive load F with simple supports at b and c. Given asimplified minimum weight structure shown in Fig. 2 withspan = L,applied force =,F2fF,where, f is the stress acting on the tension members at any point during loading.

Given asimplified minimum weight structure shown in Fig. 2 withspan = L,applied force =,F2fF,where, f is the stress acting on the tension members at any point during loading. Table 1 showsthe load relationships existing in the structure members. Structures designed based on maximumstiffness criterion must satisfy the following strain ratio (Dewhurst, 2005)2121CTTCCCTTCTEEEE(i)Inthis work, simple dual-material minimum weight structures consisting ofMetPreg®/aluminum alloy 3003 composite as tension members and aluminum alloy 3003 ascompression members were fabricated. Samples with titanium/aluminum alloy 3003 andaluminum alloy 3003 as separate structural members were also fabricated.

Samples with titanium/aluminum alloy 3003 andaluminum alloy 3003 as separate structural members were also fabricated. Both maximumstrength and maximum stiffness design criteria were used for the design of the fabricatedstructuresFigure 2: Free body diagram of the dual-material minimum weight structureTable 1: Load and Size Relationship for Minimum Weight Structure ElementLengthForceCross-sectional areaoc2/L2/2/fF)2/(CFoa2/L2fF)/(CFac2/LfF2/)2/(TF2Experimental WorkA Solidica FormationTM ultrasonic consolidation machine was used for all thefabrications in this work. During typical operation, the machine uses an automatic foil feedingmechanism, but foil materials can be fed manually when it is necessary. The parts were made onaluminum alloy 3003-H18 substrate materials of 355 x 355 x 12 mm size, mounted on a heatplate. Foil materials of aluminum alloy 3003-H18, MetPreg® and CP titanium were used for thefabricated structures.

Foil materials of aluminum alloy 3003-H18, MetPreg® and CP titanium were used for thefabricated structures. The structures are composed of tension and compression members. Thetension members carry simple tensile loads while the compression members carry simplecompressive loads when a three-point load is applied as illustrated in Fig. 2. One set of thestructures consist of MetPreg®/Al3003 composite tension members and Al3003 compressionmembers. The other set was made of titanium/Al3003 composite tension members and Al3003compression members. Both maximum stiffness and maximum strength minimum weightstructure design criteria were used for each material combination. Thus, for MetPreg®/Al3003composite and Al3003 material combination, the two criteria were used to design structureshaving different member sizes. The same criteria were applied for the titanium/Al3003 andAl3003 material combination.

The same criteria were applied for the titanium/Al3003 andAl3003 material combination. Three structures were fabricated using each criteria and materialcombination.A third set of structures were ultrasonically consolidated exclusively with Al 3003-H18foil material using the dimensions of the MetPreg®/Al 3003 material minimum weight structures.Another set of structures were fabricated using wrought Al 3003 H-18, having the same basematerial as the Al 3003 foil used. These last sets of structures were fabricated as single materialcopies of the shape and sizes of the MetPreg® and titanium reinforced dual-material structuresdescribed above. All the single material structures were fabricated for the sole purpose ofcomparing their load carrying capabilities with those of the dual-material structures.

All the single material structures were fabricated for the sole purpose ofcomparing their load carrying capabilities with those of the dual-material structures. The majorcomparison factor is the strain energy densities of the structures at failure.For the purpose of analysis and discussion in this work, the structures have been namedas follows. All structures designed based upon maximum strength design criterion or singlematerial copies of such designs have their labels hyphenated with “STR”. Similarly, thosedesigned based upon maximum stiffness criterion or their single material copies have their labelshyphenated with “STF”. Thus, MetPreg®/Al 3003 dual-material structures designed based onmaximum strength criterion are labeled Met-STR, while structures of the same materialsdesigned based on maximum stiffness criterion are labeled Met-STF. Corresponding structuresdesigned based on Ti/Al 3003 materials are labeled Ti-STR and Ti-STF.

Corresponding structuresdesigned based on Ti/Al 3003 materials are labeled Ti-STR and Ti-STF. The single materialdirect copies of Met-STR and Met-STF structures ultrasonically consolidated using Al 3003 foilsare correspondingly labeled Al-STR and Al-STF. Also, those machined out directly from Al3003-H18 plate as single material structure copies of Met-STR and Met-STF are correspondinglylabeled W-Al-MSTR and W-Al-MSTF, while copies of Ti-STR and Ti-STF are respectivelylabeled W-Al-TSTR and W-Al-TSTF.

This becamenecessary as the UC machine needed to operate in an unconventional sequence characterized byinterruptions while changing from one file to the other because of the different materials used.The first code was for consolidating the Al 3003 matrix material of 100 x 105 x 0.3mmconsisting of two layers of foils. After the consolidation, the integrated 3-axis CNC millingfacility was used to machine out the channels for accommodating the embedded reinforcingmaterials. This was used to accommodate the reinforcing materials for the respective tensionmembers in the fabricated structures. The MetPreg® foil was embedded in 9.5mm wide cavitieswhile that for the titanium foils was 12 mm.

The MetPreg® foil was embedded in 9.5mm wide cavitieswhile that for the titanium foils was 12 mm. This first step is illustrated in Fig.3a.(a): Cavity machined into deposited Al 3003 matrix(b): Reinforcing foil placed in position(c): Titanium foil of equal width with the sonotrode placed on top of reinforcing foil preparatory to indirect welding(d): The first reinforcing foil fully welded into the Al 3003 matrix(e): The second reinforcing foil fully welded into the Al 3003 matrix(f): The structure profile machined using UC integrated CNC milling(g): The structure removed from the substrate using conventional milllingFigure 3: Structure fabrication sequenceThe second and third machine codes were used to weld the reinforcing material that issandwiched with Al 3003 foils in alternate layers to make the composite tension members ineach of the structures.

In the respective structures, the reinforcing materials serve to reinforce theAl 3003 matrix foils in the tension members. The composite reinforcing foils were put in placeone at a time and welded indirectly by placing a 25mm width titanium foil between it and thewelding sonotrode as shown in Figs. 3b to 3e. The indirect welding was to prevent the sonotrodefrom having direct contact with the softer Al 3003 matrix material because of the required highwelding amplitude applied for the reinforcing materials. Direct welding can destroy the Al 3003matrix material at the high welding amplitude. The structures’ constituent materials were weldedusing different sonotrode vibration amplitudes as determined in earlier work by Obielodan et al.,2010B. Table 3 shows the compositional, hardness and dimensional details of the materials usedwhile Table 4 shows the welding parameters applied for their consolidation.

Material Vol. Fraction (%)Stiffness (GPa)Tensile strength Density (Kg/m3) (MPa)MetPreg® /Al3003 Composite661295003020Ti/Al3003 Composite2577.12322934Al3003682002730The cycle of operations described above were repeated until the desired final thickness ofthe structure was attained. Thereafter, the fourth machine code file was used to cut out the profileof the structure as shown in Fig. 3f in reverse order from top to the bottom using the integratedCNC milling head. The completed structure is shown in Fig. 3g. The fabricated structures weresubjected to three-point loading using a short beam shear test fixture (ASTM D 2344) asillustrated in Fig. 4.

The completed structure is shown in Fig. 3g. The fabricated structures weresubjected to three-point loading using a short beam shear test fixture (ASTM D 2344) asillustrated in Fig. 4. A 50kN capacity Tinius Olsen tension testing machine was used to apply acompressive load at 0.5mm/min speed until the structure failed.Figure 4: Structure under test using a 3-point bend test fixture2.1 Metallographic StudiesSmall samples cut from the intersecting joints of the composite members as well as attheir joints with the matrix material at the top of the structures were mounted and polishedaccording to standard metallographic procedures. They were observed under optical microscope.Fractographic studies were also carried out on fractured surfaces of the structures.3Results and discussion3.1 MicrostructuresMicrographs of section a-a in Fig. 3g at the intersection joints of representative structuresare shown in Fig. 5 below.

3g at the intersection joints of representative structuresare shown in Fig. 5 below. Figure 5a shows the side and end views of reinforcing MetPreg® foilsfor the left and right hand side tension members respectively of the structure shown in Fig. 3g.MetPreg® foils for the left hand side tension member have their foils stretching through thelength of the member in alternate layers. At the other layers, shorter reinforcing foils are seenwith abutting joints with those of the right hand side tension member (that is, those with their endviews shown). Each of the reinforcing foil layers are alternated with the Al 3003 matrix material.The abutting reinforcing foil joints have gaps of varying sizes, because, they were manually laidwithout any tacking operation by the sonotrode. There is some measure of foil displacement ofun-tacked foils during the welding operation resulting in the shift of positions. For structuresreinforced with titanium foils, Fig.

There is some measure of foil displacement ofun-tacked foils during the welding operation resulting in the shift of positions. For structuresreinforced with titanium foils, Fig. 5b shows corresponding foil arrangement at the intersectionjoint between the two tension members.(a): A view of interlocking MetPreg® foils in Al 3003 matrix at the intersection joints(a): A view of interlocking titanium foils in Al 3003 matrix at the intersection jointsFigure 5: Micrographs of the interlocking foils at the intersection joints of reinforced structures. The side viewed and end viewed foils belong to the left and right tension members of Fig. 3g respectively3.2 Failure StrengthsTables 6 and 7 shows the failure loads for each of the fabricated structures. It can beobserved that the structures designed using maximum strength criterion failed at higher loadswhen compared to those fabricated using maximum stiffness criterion.

It can beobserved that the structures designed using maximum strength criterion failed at higher loadswhen compared to those fabricated using maximum stiffness criterion. The failure load data canhowever, not be used for direct comparison since the structures were not exactly of the samethickness. The more useful data based on calculated strain energy density at the point of failure ispresented in Tables 8 and 9. They are also shown graphically in Fig. 6 and Fig. 7. The strainenergy density values were calculated usingVvU)21((ii)Where,U = strain energy density for the structureσ = stress in each member at the point of failureε = strain in each member at the point of failurev = volume of each memberV = total structure volumeThe stress σ in each member at the point of failure was calculated by normalizing theresolved load (based on Table 1 relationships) with respect to cross-sectional area.

With thestress obtained, the strain ε was calculated using the stiffness value of each structure member.The strain energy density data show that for structures of the same material combination,those designed based on maximum strength criterion generally have higher load carryingcapacities than those designed based on maximum stiffness criterion.

It is only in the case of Al-STR and Al-STF, ultrasonically consolidated using Al 3003 matrix material that structures basedon maximum stiffness yielded higher average strain energy density.Table 6: Failure Load (N) Data for MetPreg®/Al 3003 Based StructuresSample Met-STR1 51902 47603 4140Met-STF406050204720Ti-STR389037604120Ti-STF3210 33603500Al-STR3120 30602530Al-STF4010 38403760Table 7: Failure Load (N) Data for Ti/Al 3003 Based StructuresSample Ti-STR1 38902 37603 4120Ti-STF3210 33603500W-Al-STR320030002730W-Al-STR234022302320Table 8: Strain Energy Density (N/mm2) Data for MetPreg® Based StructuresSample Met-STR11.72E+052 1.67E+0531.79E+05Average 172585.4Met-STF1.208E+51.45E+051.43E+05136266.67Al-STR9.77E+04 8.02E+0468289.0682068.39Al-STF1.07E+05 8.64E+049.93E+0497500W-Al-STR1.22E+51.11E+51.12E+5115062W-Al-STF1.01E+59.69E+49.49E+497583Strain Energy Density Comparison for Structures Designed Using MetPreg®/Al 3003 Material Properties2.00E+05)m q s / N1.80E+051.60E+05( y t i s n e D y g r e n E n a r t Si1.40E+051.20E+051.00E+058.00E+046.00E+044.00E+042.00E+040.00E+00Met-STR Met-STFAl-STRAl-STF W-Al-MSTR W-Al-MSTFFigure 6: Strain energy densities for structures designed based on MetPreg®/Al 3003 material propertiesTable 9: Strain Energy Density (N/mm2) Data for Titanium Based StructuresSample Ti-STR11.48E+052 1.38E+0531.50E+05Average 1.45E+05Ti-STF1.48E+05 1.28E+051.13E+05129621.8W-Al-STR 8.16E+04 8.18E+047.75E+480288W-Al-STF 6.27E+04 6.83E+047.64E+0469132Strain Energy Density Comparison for Structures Designed Using Ti/Al 3003 Material Properties1.60E+05)m q S / N1.40E+05( y t i s n e D y g r e n E n a r t Si1.20E+051.00E+058.00E+046.00E+044.00E+042.00E+040.00E+00Ti-STRTi-STFW-Al-TSTRW-Al-TSTFFigure 7: Strain energy densities for structures designed based on Ti/Al 3003 material properties3.3Statistical Analysis of the Strain Energy DensityThe results of the strain energy densities presented above were analyzed statisticallyusing SAS 9.1 to verify whether or not their differences are significant.

The experiment was atwo way factorial design with three replicates. The analyses combine the results of theMetPreg®/Al3003 and Ti/Al3003 based structures all in one. Structure design criteria andmaterial used were the two fixed factors used for the analyses. For the purpose of this statisticalanalysis, alphabetical letters were assigned to groups of experimental units based on the materialused as follows: A = Met-STR and Met-STF; B = Ti-STR and Ti-STF; C = Al-STR and Al-STF;D = W-Al-MSTR and W-Al-MSTF; and E = W-Al-TSTR and W-Al-TSTF. Thus, material as afactor comprise five levels while the design as a factor comprise two levels 1 and 2corresponding to structures designed based on maximum strength and maximum stiffnessrespectively.The result of the analysis shows that the data satisfies the assumption of approximatenormality and homoscedasticity.

Analysis of variance (ANOVA) in Table 10 shows that the twofactors and their interactions have significant effects on the response variable. Using factor-levelcomparison in the least square means (LSMEANS) table (that is, Table 11 below) and theinteraction plots in Fig. 8, it is evident that the difference of means between Met-STR and Met-STF is statistically significant. Met-STR and Ti-STF also have significantly different means.However, the difference between the means of Met-STR and Ti-STR is marginally insignificant.Between Ti-STR and Ti-STF as well as between Met-STF and Ti-STF there is no significantdifference of mean strain energy density.Comparing each of the reinforced structures with the un-reinforced ones, Met-STF andW-Al-MSTR have statistically insignificant different means, although the former yielded higheraverage strain energy density.

Also, Ti-STF and W-Al-MSTR have statistically insignificantdifferent means, although the former yielded higher average strain energy density. Apart fromthese two cases, all reinforced structures have significantly different mean strain energy densitieswhen compared with the single material structures. Post hoc means analysis data in Table 12 forthe individual factors shows that structures designed based upon maximum strength criterionhave significantly higher average strain energy density than those designed based uponmaximum stiffness criterion. Also, all the material categories have statistically significantdifferent average strain energy densities with MetPreg® reinforced materials yielding the highestvalue.From this analysis, it can be inferred that UC fabricated structures with appropriatereinforcement leads to significant improvement of their load carrying capability compared tofabrications with the matrix materials only.

Although MetPreg® reinforced structures performedbetter than titanium reinforced ones with corresponding design criteria, Ti can be considered as agood reinforcement material. This is because the Ti volume fraction of 25% was considerablylower than the 60% volume fraction of MetPreg® in the same matrix material. The higher cost ofMetPreg® makes titanium a good alternative, although the former has higher specific strengththan the latter.Table 10: Analysis of Variance of the Experimental DataTable 11: Tukey’s LSMEAN Adjustment for Multiple Comparison(a)(b)Figure 8: Interaction plots between the factorsTable 12: Post Hoc Means Analysis for the Individual Factors12a : Design12b: Material3.4Failure FeaturesThe failure features of the structures depend mostly on the design criterion used. Most ofthose designed based on maximum stiffness criterion failed at the flange as shown in Fig. 9 forany material combination.

Most ofthose designed based on maximum stiffness criterion failed at the flange as shown in Fig. 9 forany material combination. Their flange widths are generally smaller than those designed basedon maximum strength criterion. It shows that higher stresses are concentrated at the neck of theflanges. Rather than fracture by tearing the materials, most of them deform and in some cases,the consolidated foils delaminate as shown Fig. 10. However, none of the maximum strengthstructures failed at the flange. MetPreg® reinforced maximum strength structures generally failedat the left hand side tension members. The failures occurred on those members at the edge-to-edge foil joints of the Al 3003 matrix materials or at the edge-to-edge joint of the reinforcingMetPreg® materials. The right hand side tension members did not have any foil joint; as such nofracture occurred on them.

The right hand side tension members did not have any foil joint; as such nofracture occurred on them. The left tension members were cut perpendicular to the direction ofconsolidated foils, this make them to have intra-layer edge-to-edge joints. The right hand tensionmembers were however, cut along the direction of foil consolidation. The properties of the jointshave been characterized in earlier work (Obielodan et al., 2010A)Figure 9: Failure feature of a structure designed based on maximum stiffness criterionFigure 10: Delamination at the flange for some of the structures designed based on the maximum stiffness criterionFigures 11 to 13 illustrate the modes of failure for the MetPreg®/Al3003 reinforcedmaterials. In Fig. 11, failure occurred at the foil edge-to-edge joints of the matrix material. Thestress on the tension member for this structure at failure was 313MPa.

In Fig. 11, failure occurred at the foil edge-to-edge joints of the matrix material. Thestress on the tension member for this structure at failure was 313MPa. A defective matrixmaterial foil joint must have exposed the reinforcing MetPreg® foil to cause failure. Figure 12shows a combination of failures at the matrix foil joints and the joints of the reinforcingmaterials. There were inter-lamina foil delaminations between the fracture locations. The stresson the tension member at the point of fracture was 320MPa. In Fig. 13, failure occurred at theedge-to-edge joint of the reinforcing MetPreg® foils only. The tension member failed at 309MPa.It is worthy of note that most ultrasonically consolidated MetPreg®/Al3003 composite (ofthe same MetPreg® volume fraction, that is 60%) tensile specimens preliminarily tested beforethese structures were fabricated failed prematurely at stresses ranging from 300 to 450MPa.

TheAl2O3 reinforcing fibers in the MetPreg® foils may have been damaged under the action of theultrasonic energy applied through the sonotrode during welding, resulting in premature brittlefailures of the MetPreg®/Al3003 composite members.

For the structures designed based onmaximum strength criteria using the MetPreg®/Al3003 material combination, the stresses on thetension members with foil edge-to-edge joints generally determined the failure point.Figure 11: Fracture at the edge-to-edge foil joint of the matrix material on a MetPreg® reinforced tension member.Figure 12: Fracture at both the edge-to-edge foil joint of the matrix material and the edge- to-edge joint of the reinforcing foils on a MetPreg® reinforced tension member.Figure 13: Fracture at the edge-to-edge joint of the reinforcing foils on a MetPreg® reinforced tension member.Details of the failure stresses on the tension members of structures designed withMetPreg®/Al3003 using maximum strength criterion and their single material copies are shownin Fig. 14 (that is, those described with STR labels).

14 (that is, those described with STR labels). Al-STR structures ultrasonicallyconsolidated exclusively with Al3003 foils and default foil overlap setting of width 23.90mmfractured at the tension member at an average stress of 169MPa. This fracture stress is within therange of tensile strengths obtained for tensile specimens fabricated with the same machineparameters in earlier work (Obielodan et al., 2010A).

This fracture stress is within therange of tensile strengths obtained for tensile specimens fabricated with the same machineparameters in earlier work (Obielodan et al., 2010A). Also, the W-Al-MSTR structuresfabricated exclusively with the wrought Al3003 substrate material failed at the tension memberat an average stress of 205MPa, which is within the range of the tensile strengths obtained for theparent material.Stress on the Tension Members of Structures Designed Using MetPreg/Al3003 Material Properties350300) a p M250(s s e r t s el i s n e T200150100500Met-STR Met-STFAl-STRAl-STF W-Al-MSTR W-Al-MSTFFigure 14: Stress on the tension members of structures designed using MetPreg®/Al 3003 material properties at the point of failure4ConclusionsThe use of the ultrasonic consolidation process for the fabrication of multi-materialminimum weight structure has been demonstrated.

A fabrication methodology for joining foils ofdifferent materials to make dual-material structures was developed. Test results show that thereare significantly higher strain energy densities at the point of failure in structures with reinforcedmembers compared to those fabricated with Al 3003 matrix materials only. As a result, their loadcarrying capacities were greatly improved. It was observed that the failure mode of the structuresis generally dependent upon the design criteria and the materials used. Structures fabricatedbased upon maximum strength criterion using MetPreg®/Al3003 composite materials as thetension members exhibited brittle failures at edge-to-edge foil joints. Those designed based uponmaximum stiffness criterion generally failed at the flange of the triangular structure irrespectiveof the material combination.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDynamic Build Bed for Additive ManufacturingIsmail Enes Yigit and Ismail LazogluManufacturing and Automation Research Center, Koc University, Mechanical Engineering Department, Sariyer, Istanbul 34450, TurkeyAbstractCompared to subtractive manufacturing, additive manufacturing generally has low material waste. However, models with large overhangs require manufacturing of support structures which ends up as waste material. This paper proposes the use of a dynamic build bed for reducing support structures. The bed consists of an array of actuated pins which move in the build orientation. Each pin can be individually moved to the correct height for supporting the given model. Two separate applications of the build bed are investigated.

Each pin can be individually moved to the correct height for supporting the given model. Two separate applications of the build bed are investigated. In the first application, the dynamic build bed is used as support structures in deposition-based AM methods. The pins individually raise out of the build bed to support the overhang geometry at the given deposition height. The second application is in powder-based AM methods. In the second application, the pins are used to fill the space of the powder where the geometry will not occupy. The pins are individually lowered in the build orientation to make space for a new powder layer. Thus, saving excessive deposition of powder.Keywords: Support structures, Dynamic, Additive Manufacturing1. IntroductionAdditive manufacturing (AM) is praised to have low material waste in comparison to traditional manufacturing methods such as subtractive manufacturing. This is not valid for all given computer aided design (CAD) models.

This is not valid for all given computer aided design (CAD) models. In deposition-based AM methods, the material needs to be deposited on an existing surface. When the given CAD model has overhangs, the material cannot be directly deposited. In this case the layer of the model needs to be supported by support structures. Support structures are manufactured just as the model itself only to support the overhang geometries. These support structures are not part of the final CAD geometry. Thus, they need to be manually removed after manufacturing and they end up as waste material. These support structures also increase the manufacturing and the require manual labor time.In powder-based AM methods, the overhangs do not cause trouble since the powder is equally deposited to cover the whole layer at a time. This powder fills the space under the model and acts as a natural support structure for the whole geometry. Thus, any overhang of the model is also supported by the deposited powder.

This powder fills the space under the model and acts as a natural support structure for the whole geometry. Thus, any overhang of the model is also supported by the deposited powder. However, the powder degrades once after it has been through a manufacturing cycle and becomes waste. Therefore, it is beneficial to deposit the minimum amount of powder on each layer. This paper proposes the use of a dynamic build bed for reducing the manufacturing of support structures in the first application and reducing the deposition of powder material in the second application.The bed consists of an array of actuated pins which move in the build orientation. Each pin can be individually moved to the required height for supporting the given model. The proposed dynamic bed system has been previously utilized in human computer interaction (HCI) field, but it has not been utilized in AM.

The proposed dynamic bed system has been previously utilized in human computer interaction (HCI) field, but it has not been utilized in AM. Sean Follmer and Daniel Leithinger have introduced the1568inFORM system which is a 2.5D shape display consisting of 30x30 array of linearly actuated pins [1] given in Figure 1. Each pin can dynamically move in the Z orientation for interaction with the user or manipulation of the physical objects on the display. A single pin can exert 1 Newton of force and extend up to 100mm from the surface. The height map of the pins is calculated by OpenGL shaders and are sent to the 900 pins via a microcontroller.P n siL i n k a g e sA c t u a t o r sFigure 1: The inFORM system with 900 mechanical actuators connected to the pins with the linkages [1]Leithernger et. al later used the inFORM shape display to physically render shapes and geometries [2].

al later used the inFORM shape display to physically render shapes and geometries [2]. They also proposed that users can use these shape displays to quickly render physical CAD models before they are additively manufactured. Rendering before AM will improve the prototyping cycles. In this paper, it is proposed to use the given physical renders as support structures in the AM process.Several work in the literature proposes methods to reduce support materials. Some of these works are focused on multidirectional manufacturing where the deposition tool has more than three degrees of motion and can deposit in non-conventional orientations [3]. Another method is placement of foreign object as support structures in convenient locations where the deposition head can continue deposition without collision [4].

Another method is placement of foreign object as support structures in convenient locations where the deposition head can continue deposition without collision [4]. Support structure optimizations are applied such as the sloping wall support structures [5] where the total volume of the support structure is reduced by reducing the cross sectional area of the support with the minimum self-supporting angle. Another method is support slimming to optimize the geometry by changing its shape to be more self-supported [6]. Optimization of the build orientation of the model [7, 8] is also an effective approach where the original model can be manufactured on the conventional supports at an orientation requiring the minimum amount of support.1569The proposed method in this paper has been practiced by the flexible support structures [9]. In this paper a flexible platform consisting of actuated basic units used.

In this paper a flexible platform consisting of actuated basic units used. The paper includes a control module for support structure reduction and automatic unloading. Support structure reduction mainly includes 3 steps. The first step is to discretize the model’s printing space to pieces. The second step is to identify the lowest point of each piece. The third step is to offset the base of support structure. This paper adds to the flexible support structure by introducing a universal algorithm valid for both deposition-based AM systems and for powder-based AM systems.This paper starts with introducing the conventional support structures from the literature in section 2. Then in section 3, dynamic support structures are introduced. Two separate applications of the build bed are investigated in the following sections. The first application given in section 4 is the dynamic build bed in deposition-based AM methods.

Two separate applications of the build bed are investigated in the following sections. The first application given in section 4 is the dynamic build bed in deposition-based AM methods. In this application, the pins individually raise out of the build bed to support the overhang geometries at the given deposition height. The second application, given in section 5, is the use of dynamic build bed in powder-based AM. In this application, the pins are used to fill the space of the powder where the geometry will not occupy. The pins are individually lowered in the build orientation to make space for a new powder layer. The paper ends by giving the conclusion.2. Conventional supportsAM is a process where the given CAD model is manufactured layer by layer. The CAD model is usually given as a Stereolithography (STL) file [10]. This file consists of a triangulated mesh of the model.

These regions are then sampled in the two- dimensional plane parallel to the build bed. The sampled points give the support points. The sampling quality depends on the spacing of the support structures. For each support point a ray is cast towards the build bed [11]. If the ray hits the model, the support structure starts on the model. If the ray reaches the build bed the support structure starts from the build bed and is eligible to be replaced by the pins of the dynamic build bed. Dynamic support structures are studied in the next section.3. Dynamic support structuresThe dynamic support structures consist of a two-dimensional array of pins as introduced in section 1. Each pin can move independently in the build orientation. The geometry of the pins is identical and is in the shape of a square prism with user set width (cid:1875) and height (cid:1860) as shown in Figure 1. In this figure, total of 361 pins are shown where some of them are raised for illustration.

In this figure, total of 361 pins are shown where some of them are raised for illustration. The number of pins can be set depending on the required build bed dimensions. The width and height of the pins together with the number of pins give the pin configuration.1570(1)The proposed method in this paper calculates the dynamic support structures for any CAD model with any given pin configuration.(cid:1875)(cid:1875)(cid:1860)Figure 2: Dynamic build bed of 361 pins with width (cid:1875) of 10[(cid:1865)(cid:1865)] and height (cid:1860) of 150[(cid:1865)(cid:1865)]After the pin configuration is set, the next step is to find the support points as introduced in section 2. As the width of the pins gets coarser, the pins tend to collide with the manufactured model. To deal with this, a ray is cast from each corner of the pin. Then collision is checked in the same way which the support points were checked. The results are as shown in Figure 3.

To deal with this, a ray is cast from each corner of the pin. Then collision is checked in the same way which the support points were checked. The results are as shown in Figure 3. In this figure the points where the ray hits the model are given in red. The points of the model which require support are highlighted in green. Any highlighted point which is reachable by with without colliding with the model is reached by a pin. Here the whole model shown where the pins are at their final positions. During the AM process, the pins will raise in increments of the layer height. After each layer is deposited, the pins will raise as much as the layer height to support the new layer. When a pin reaches its final height, it will be fixed at that position until the process is reset.

When a pin reaches its final height, it will be fixed at that position until the process is reset. The pins can also move individually in a wave motion to release the from the build bed for easy removal.1571Figure 3: Support points of the Stanford BunnyAs introduced previously, the dynamic build bed can be used in two different applications. The figures up to this point have all illustrated the application for deposition-based AM where the pins are raised out of the build bed. The next section will investigate the use of the dynamic build bed in deposition-based AM.4. Dynamic build bed in deposition-based AMThe first application of the dynamic build bed is the usage of the pins as support structures in deposition-based AM methods. The pins individually raise out of the build bed to support the overhang geometry at the given deposition height. At this point the pins will move either upwards or downwards depending on the AM method.

At this point the pins will move either upwards or downwards depending on the AM method. Simulations have been conducted for both applications with two different models to verify the proposed dynamic build bed method. Figure 4 and Figure 5 illustrate the results for deposition-based AM.Figure 4: Dynamic build bed in deposition-based AM with Ataturk Model1572Figure 5: Dynamic build bed in deposition-based AM with Bunny ModelThe corner with the lowest collision height is taken as the maximum height which the pin can raise. After the maximum height is calculated for each pin, these values are stored on a two- dimensional matrix. When the manufacturing has started, each pin is raised as much as the layer height when the extruder moves to the next layer. at each step, the maximum height of the pin is checked. If the pin has reached its maximum value, it is fixed at that position. This means that the pin is ready to support the model and should not move anymore.5.

If the pin has reached its maximum value, it is fixed at that position. This means that the pin is ready to support the model and should not move anymore.5. Dynamic build bed in powder-based AMThe second application is in powder-based AM methods. In powder-based methods, there is no need for special support structures as the powder itself acts as a natural support structure. In the case that most of the build box is empty, the powder that is laid to fill up the layer will go through a heat cycle and deteriorate over time. Since the powder is usually an expensive resource, it is in our best interest to lay the least amounts of powder. Here the pins are used to fill the space of the powder where the geometry will not occupy. The pins are individually lowered in the build orientation to make space for a new powder layer. Thus, excessive deposition of powder prevented.

The pins are individually lowered in the build orientation to make space for a new powder layer. Thus, excessive deposition of powder prevented. Figure 6 and Figure 7 illustrate the results for powder-based AMFigure 6: Dynamic build bed in powder-based AM Ataturk Model1573Figure 7: Dynamic build bed in powder-based AM with Bunny ModelThe manuscript is ended with the conclusion.6. ConclusionDynamic build bed for reducing the use of support structures and powder has been introduced. An array of actuated pins which move in the build orientation assist the AM process. Every pin is individually actuated to reach the desired position under the CAD model. Two separate implementations are given in this paper. The first implementation is for deposition based AM where the dynamic build bed is used for support structures. Each pin raises to support the overhang geometry at the required height.

The first implementation is for deposition based AM where the dynamic build bed is used for support structures. Each pin raises to support the overhang geometry at the required height. The second implementation is in for powder- based AM where the pins fill the space where the powder originally occupies. Each pin is lowered in the build orientation to make space for the model’s new layer. This results in less usage of powder. To verify the improvement of the proposed implementations, two models are simulated, and the results are compared to the conventional methods. Minimizing the fabrication of support structures, the material waste can be reduced, which also results in a reduction in fabrication time. Future work includes testing the proposed method on a mechanical setup.7. AcknowledgementsThis research was supported by the Koç University Manufacturing and Research Center.8.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceDynamic Defect Detection in Additively Manufactured Parts using FEA SimulationKevin Johnson1, Aimee Allen1, Jason Blough1, Andrew Barnard1, David Labyak1, Troy Hartwig2, Ben Brown2, David Soine2, Tristan Cullom3, Edward Kinzel4, Douglas Bristow3, Robert Landers31Department of Mechanical Engineering-Engineering Mechanics, Michigan Technological University, Houghton, MI 499312Honeywell Kansas City National Security Complex, Kansas City, MO 641473Department of Mechanical Engineering and Aerospace, Missouri University of Science and Technology, Rolla, MO 654014Department of Aerospace and Mechanical Engineering, University of Notre Dame, Notre Dame, IN 46556Abstract: The goal of this paper is to evaluate internal defects in additively manufactured (AM) parts using FEA simulation.

The resonant frequencies of parts are determined by the stiffness and mass involved in the mode shape at each resonant frequency. Voids in AM parts will change the stiffness and mass therefore shift the resonant frequencies from nominal. This paper will investigate the use of FEA to determine how much a void size, shape, and location will change the resonant frequencies. Along with where the optimal input and response locations are in order to find these frequency changes. The AM part evaluated in this work includes a common tensile bar and hammer shaped part evaluated individually and as a set of parts that are still attached to the build plate. This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1. IntroductionAdditive manufacturing continues to be researched as a viable method for making parts.

IntroductionAdditive manufacturing continues to be researched as a viable method for making parts. Some of the potential advantages are the capability to make complicated shapes and less wasted energy due to scrap of material that went through an extensive process to create. But currently a significant negative is the relative long time to manufacture compared to other manufacturing processes. In addition to this manufacturing time is time to test for part voids. Voids in parts negatively affect a parts strength and ductility. Very accurate testing methods like x-ray and CT- scanning are successfully being used to find voids [7]. But these methods are slow and inexpensive.Research in dynamic testing has been done to find inexpensive and relatively fast ways to test parts for voids [1]. Dynamic testing is a process where a part is dynamically excited and the1281response is measured. The excitation is usually applied with a shaker or with a modal hammer.

Dynamic testing is a process where a part is dynamically excited and the1281response is measured. The excitation is usually applied with a shaker or with a modal hammer. The input signal into the test part by a shaker is measured with an accelerometer on the base plate of the shaker. The input signal into the test part by the modal hammer is measured by a load cell in the tip of the hammer. The response is typically measured with an accelerometer. The Fourier Transform is applied to these signals to take them into the frequency domain. Much like the Fourier Transform is applied to sound to extract all of the different frequencies that add up to create the actual recorded sound, the response signal of a part or structure is taken through the Fourier Transform to extract all of the resonant frequencies and mode shapes. The response signal is then normalized to the input signal by dividing it by the input signal and plotted against frequency.

The response signal is then normalized to the input signal by dividing it by the input signal and plotted against frequency. This is called a Frequency Response Function (FRF). The amplitude of the FRF is usually shown in the log scale to help capture the small resonant frequencies. The peaks of the curve are resonant frequencies where the response per unit input is large. The dips are anti- resonances where the part or structure doesn’t respond to the input. The mode shapes are how the part or structure moves at each resonant frequency. Each resonant frequency is determined by the stiffness (k) and mass (m) involved in the mode shape at that frequency as shown in Equation 1.(1)(cid:2033) = (cid:3496)(cid:1863) (cid:1865)If the stiffness or mass of a part or structure is changed by a void or change in density the resonant frequencies with affected mass or stiffness that participates in the mode shape at that frequency will change.

These frequencies with a lot of mass participation or stiffness change can be used to find the faults. This is where Finite Element Analysis can help with the physical testing. Each particular build plate of parts or individual parts can be analyzed with FEA to determine the resonant frequencies. This can be done up to the highest frequency that can be found with the testing method available. The frequencies with the highest difference with a nominal void-free part can be quickly determined. Then FEA can be used to apply the excitation and measure the response at different locations to determine the optimal locations for testing. The testing time will be shortened significantly. This article demonstrates the use of FEA to assist in physical dynamic testing of different AM builds. Before the builds in this article were analyzed with FEA test locations were randomly decided based on what the mode shapes might look like.

Before the builds in this article were analyzed with FEA test locations were randomly decided based on what the mode shapes might look like. FEA was then done to assist in finding the parts with voids that weren’t found with the first round of testing. As seen below if FEA was done in the beginning it would have eliminated extra rounds of testing.2. MethodsThe Finite Element Analysis was done using HyperMesh by Altair [10] and SolidWorks Simulation [11]. All parts were meshed using a mixture of brick and second order tetrahedral elements and also with midplane shell elements. The mesh consisted of a minimum of four elements across a thickness (cid:82)(cid:73)(cid:3)(cid:167) 0.095 mm. The resonant frequencies and mode shapes were extracted using the EIGRL Card. Experimental impact data from a hammer test was recorded in1282the time domain and imported as a table in Hypermesh for the impact signature. An impact force of 1 N was induced into the part.

Experimental impact data from a hammer test was recorded in1282the time domain and imported as a table in Hypermesh for the impact signature. An impact force of 1 N was induced into the part. Responses were recorded at different nodes using the Global Output Request.For FEA the equilibrium matrix equation of motion for a linear elastic material is as shown:[(cid:1839)](cid:3427)(cid:1847)(cid:4663) (cid:3431) + [(cid:1829)](cid:3427)(cid:1847)(cid:4662) (cid:3431) + [(cid:1837)][(cid:1847)] = [(cid:1832)]Where [M] is the mass matrix, [C] is the damping matrix, [K] is the stiffness matrix, and [U] is the displacement. For extracting the modes damping is ignored and assume no external forces.

These builds were printed with the Selective Laser Melting Process (SLM). The material properties of the bars found experimentally are Young's Modulus 180 GPa, density 7.85 g/cm3, and Poisson's ratio 0.24. The first build consisted of nine groups of randomly placed tensile bars on the build plate. One group, the nominal group, had no intentional defect added. The remaining eight groups had defects that consisted of powder that was intentionally not melted in the center of the part as shown in Figure 1b. The voids in these eight groups all have the same cross section as shown in Figure 1b and vary in height from 50 μm to 400 μm.

The voids in these eight groups all have the same cross section as shown in Figure 1b and vary in height from 50 μm to 400 μm. The void leaves a 0.020” (0.50mm) solid perimeter on the outside of the bar.1283(2)(3)(4)(5)(6).02in[0.50mm]TYP I II [3.81 mm].15inIJDIIJI ~ [ 6 .35 mm]I.25in.11.2150µm to 400µmyl_ (l1 .75zX(1a)(1b)Figure 1: Tensile Bar build provided by MS&T (1a), Tensile Bar detail showing layer defect location (2a)The first 20 modes of a all 5 groups found with FEA are shown in Table 1. These were analyzed individually with a fixed constraint on the base of the bars. Column 2 also has the average of the groups of bars of some modes found experimentally. The resonant frequencies found experimentally are within 2% the FEA results. The main takeaway from the FEA results are that the difference in frequencies of all of the tensile bars with voids from the nominal tensile bars are very small.

The main takeaway from the FEA results are that the difference in frequencies of all of the tensile bars with voids from the nominal tensile bars are very small. The first bending mode in the X direction is less than one hertz different (0.5 Hz to 0.79 Hz). This is due to the stiffness in bending mainly affected by the outside fiber of the material and the void is in the center of the tensile bars. During experimental testing of the tensile bars the frequency resolution needs to be set low enough to distinguish between these small frequency differences. Table 1 also shows the mode with the largest frequency change is the ninth mode. This is the first axial mode of the tensile bar. The entire cross-sectional area affects the stiffness of the bar in the axial direction. The axial mode should be targeted during experimental testing for the best chance at finding the parts with voids.

This would make it difficult to capture the x-direction modes which FEA has found to be better. Figure 3 shows the FRF’s with the input in the z-direction and responses in the z-direction and x-direction. If the response is only measured in the z-direction the x-direction modes are obviously missed. As expected the response in the x-direction is many orders of magnitude lower than the z-direction. Especially the center response as the torsional movement is not recorded. Of course, if the x-direction modes are desired it would be best to excite in the x- direction.l .0Er011.oe•OO1.0E-011 OE-02Z-Dir:Z-Dir: Corner Input Center Input Center Res Center ResZ-Dir: C Input CoZ-Dir: Corner Input Corner Res1.0E-03i E1 OE-041.0E--05Z-Dir Center Input ,I X-Dir Corner Res1.0E-061.0E-07,.oe..osZ-Dir Center Input / X-Dir Center Res1.0E-0&-062512503 1253150 437!> F reque ncy (Hz)600066256250..

,.150<>8 125I/ { (l(~Figure 3: FRF’s of Nominal Tensile Bar Fixed at BaseExperimentally testing the tensile bars on the build plate comes with other difficulties. Thefirst axial mode of the tensile bars at 9755 Hz was found to be the best mode for testing for voids. This is the mode of the tensile bar analyzed individually and not attached to a build plate. When added to a build plate with other tensile bars the mass of the build plate is involved in the modes. This causes the tensile bars axial modes to shift to other frequencies and not all tensile bars are involved at each frequency as can be seen in Figure 4. The base plate also has complicated modes around the same frequency that will also be moving the bars in the axial direction. These were analyzed with fixed translational constraints at the center of the mounting holes.

These were analyzed with fixed translational constraints at the center of the mounting holes. In order to excite the bars without the base plate the base plate would have to be suppressed or dampened.1286(4a)(4b)(4c)(4d)Figure 4: Modes at 9,395 Hz (4a), 9,451 Hz (4b), 9,558 Hz (4c), and 9,947 Hz (4d)3.2 Hammer Shaped Test PartsThe next chosen part for testing was a hammer shaped profile shown in Figure 5. This shape was chosen to give a number of good measurable modes in the frequency range up to 8,000 Hz for testing purposes. The parts were printed on 3D Systems ProJet MJP 3600 owned by Michigan Tech University with VisiJet M3-X ABS like material. Each build was printed with 4 tensile bars. Physical testing showed the parts to have a Modulus of Elasticity of approximately 2,000 MPa which is slightly under the 2,168 MPa published by 3D Systems.

Each build was printed with 4 tensile bars. Physical testing showed the parts to have a Modulus of Elasticity of approximately 2,000 MPa which is slightly under the 2,168 MPa published by 3D Systems. The published density of 1.04 g/cm3 was used.12872.50 1y---r30°TYPzX .75 ~.75Figure 5: Test PartTest parts were analyzed with five different configurations shown in Figure 6; no void (nominal), void 1, void 2, void 3, and void 4. The voids were spherical in shape with a diameter of Ø0.25” and placed in locations where theoretically different frequencies/mode shapes would be needed to find the differences with the nominal part.No VoidNo VoidVoid 1Void 2Void 3Void 4Figure 6: Test Parts with voidsThe first 27 modes of all test parts found using FEA are shown in Table 2. These wereanalyzed individually with a fixed constraint on the base of the bars. All parts were meshed using second order tetrahedral elements.

These wereanalyzed individually with a fixed constraint on the base of the bars. All parts were meshed using second order tetrahedral elements. The mesh consisted of a minimum of four elements across a thickness of (cid:167) 5.4 mm. Modes with high percent difference and/or the largest difference in frequency with the nominal part are highlighted in yellow. These are the frequencies that have the highest chance of being able to distinguish the parts with voids from the nominal parts.1288No.

The mode shape for this frequency is the first order bending mostly about the z-axis as shown in Figure 7a. The second mode is a bending mostly around the x-axis with a frequency of ~115 Hz as shown in Figure 7b. These two close modes can be difficult to differentiate depending on how the part is excited and where the response is measured.zX(7a)(7b)Figure 7: Test Part mode 1 (7a), mode 2 (7b)1289If the impact and response locations were initially chosen before FEA as shown in Figure 8a,with the resulting FRF shown in 7b. It is easy to realize that the x-direction response will be more at the first mode and the z-direction response will be more at the second mode. The 0.88 Hz shift at mode 1 and the 0.57 Hz shift at mode 2 for void 1 can be seen. The interesting thing to notice is that the void 1 part shows more x-direction response at mode 2 than the other bars.

The interesting thing to notice is that the void 1 part shows more x-direction response at mode 2 than the other bars. This is because the void in void 1 part is at a location that it changes the direction of mode 2 enough to have more response in the x-direction than the other bars. Not only is there a shift in frequency, there is also a change in the mode shape translation direction caused by the void.Point 6 Response _ FEA1.00E+o0ImpactAll Others X-Dir\y.; ~ 1.00E-01Void 1 X-Dir has more response at mode 2 than the other parts.zI.O0E-02100105110Frf!nmm r:.v l H:z)115120(8a)(8b)Figure 8: Test Part 1st 2 ModesAs can be seen in Table 2 other modes for detecting void 1 other than mode 1 are modes 17 and 26. These are at ~4404 Hz and ~7528 Hz respectively.

These are at ~4404 Hz and ~7528 Hz respectively. The mode shapes for these modes can be seen in Figure 9.(9a)(9b)Figure 9: Mode 17 - 4404Hz (9a), Mode 26 - 7528 Hz (9b)1290125If the excitation and response locations were chosen to be as shown in Figure 10a the response for mode 17 would be as shown in Figure 10b. As shown in Figure 10b the response amplitude is very low. This is due to the test part being excited in a direction there is basically no response for mode 17. It can be seen though that void 3 and void 4 parts do have some response.

This is due to the test part being excited in a direction there is basically no response for mode 17. It can be seen though that void 3 and void 4 parts do have some response. This is again due to the fact that the void in void 3 and void 4 parts are in a location that effects the mode shape direction at the impact and response locations.Location 1 Response Mode 17 - FEAImpact on sideLocation 1 1 on endLow response magnitudeNo Void Y-Di - Void 1 Y-Dir - Void 2 Y-Dir - Void 3 Y-Dir - Void 4 Y-DirzX1.00E-05430043204340436043804400 Frequency (Hz)44204440446044804500(10a)(10b)Figure 10: Excitation and Response Locations (10a), FRF’s (10b)FEA results show better locations for the excitation and response and can be seen inFigure 11. All parts have a measurable response and frequency shifts from the nominal. Void 1 part has the most shift with void 3 and void 4 parts having more than void 2.

All parts have a measurable response and frequency shifts from the nominal. Void 1 part has the most shift with void 3 and void 4 parts having more than void 2. Void 2 has almost no mass participation in this mode shape. Mode 18 is only 15 Hz higher than Mode 17 but doesn’t show up in the FRF. This is because the impact at location 7 and response at location 1 don’t have a significant amount of participate in this mode shape. The locations of the excitation and response can be strategically placed to excite only the modes desired and have other modes not excited.1291Location 1 Response Mode 17 Void 1 - FEALoe 1 on endVoid 3Mode 17 -4400 Hz.:1 .50E-0:d43504 '!604 '1704 3804 390Frequency (Hz)440044 10y4420'\ Bad impact for Mode 18XMode 18 -4415 HzFigure 11: FRF’s with Excitation at Location 7The best modes for finding Void 2 were found to be mode 7 and mode 12 as shown in Figure 12.

The original excitation and response locations happen to work well for exciting and measuring mode 7.Location 6 on endImpact on side(12a)(12b)(12c)Figure 12: Mode 7 at 1192 Hz (12a), mode 12 at 2999 Hz (12b), excitation and response locations (12c)Figure 13 shows the FRF with the impact and responses shown in Figure 12. Not only does void 2 part have a significant shift from the nominal but so does void 3. Void 3 which is in the larger claw on the part has significant mass involved in this mode shape. Mode 12 is a rotational mode of the hammer head. It makes sense that there is mass involved at the location of the void in the void 2 part but it would be harder to measure and excite this mode shape. Both would have to be at the top edge of the hammer head.

Both would have to be at the top edge of the hammer head. A laser vibrometer would be able to pick up the1292translational motion of the rotation but the glue on accelerometers used covers almost half of the face height and would not record as much translational motion.Point 6 Response Void 2 Mode 7 - FEAVoid 2 - 1187 HzVoid 3 -1 188 Hz~ All other parts,.,;:::::;::::_:;:::iiloo_,/ - 1192 Hz-No Void Z-Dir Void 1 Z-Di·Void 2Z-Di"Void 3Z-D i"Void 4Z-D i"UXlE-02 ~ - - - - - - - - - - - - - - - - - - - - - -~ 1150 11 55 11 60 1165 1170 11 75 1180 1185 11110 1Hl!i 1200 1205 1210 1215 1220 Frequency (Hz)Figure 13: FRF impact on side response at Location 6Modes 13 at ~3311 Hz and mode 27 at ~7968 Hz are the best for finding the void 3 part. The mode shapes are shown in Figure 14. Mode 13 is a rotation mode on the leg with the void in void 3 so the better choice is Mode 27.

The mode shapes are shown in Figure 14. Mode 13 is a rotation mode on the leg with the void in void 3 so the better choice is Mode 27. FEA shows the original excitation locations are not the best choice for exciting this mode shape.(14a)(14b)Figure 14: Mode 13 at ~3311 Hz (14a), mode 27 at ~7968 Hz (14b)Modes 21 at ~5950 Hz and mode 27 at ~7968 Hz are the best for finding void 4. The mode shapes are shown in Figure 15. Mode 21 is a rotation mode on the leg with the void in void 4 so the better mode to use is Mode 27. FEA shows the original excitation locations are not the best choice for exciting this mode shape.1293(15a)(15b)Figure 15: Mode 21 at 5950 Hz (15a), Mode 27 at 7968 Hz (15b)A better excitation and response location for mode 27 and the FRF’s are shown in Figure 16. The frequency shifts for void 3 and void 4 parts can be seen with this mode. Void 2 has the most response per input at this mode shape. The reason for this can be seen in the mode shape.

The frequency shifts for void 3 and void 4 parts can be seen with this mode. Void 2 has the most response per input at this mode shape. The reason for this can be seen in the mode shape. The location of the void in Void 2 has the most mass involved in this mode shape. Mode 3 has the least.Void 3 & 4 Point 8 Response - FEAImpact on sideLocation 8 on side, ,Void 3 (-21 L'.f) & Void 4 (-23.13 L'.f) Shifted downNo Void Z-DirVoid 1Z-OiVoid2Z-OiVoid3Z-Oi·>) ~zX3.00E-04 ~ - - - - - - - - - - - - - - - -~ 80507QOO71l2571l5071l75 Frequency (Hz)00008025Void4Z-Oi·Figure 16: FRF and excitation/response locations for Mode 27 at 7968 Hz4. Discussion and Future Work4.1 Discussion of ResultsThis paper showed that Finite Element Analysis is a good tool for assisting in physical dynamic testing of parts to find voids. The location of the input determines if a mode shape is excited or not.

The location of the input determines if a mode shape is excited or not. If the mode shape can be measured or not is determined by the location of the response. Even the most experienced dynamic systems person might not excite or measure the modes they are looking for the first time. Significant time in retesting can be eliminated if the part can first be analyzed with FEA to determine the mode shapes and then used to determine the locations of the excitation and response.1294FEA can also determine the magnitude of the response of each void and if there is a mode shift caused by a void. As seen in the results above a voids location can change the mode shape of the part and the response amplitude per unit input.4.2 Future WorkFuture testing can look into how much the amplitude of the part changes with a void and how much the mode shape changes.

If the change in mode shape is significant enough to be measured by a full field measuring system like Digital Image Correlation it might be used along with a frequency shift to find voids in parts.Future testing can also look into how to limit the mode shapes of the build plate so the parts can remain on the build plate for testing. This would save significant time in testing. This could consist of moving the base plate modes away from the parts modes of interest.5. AcknowledgmentsThis work was funded by the Department of Energy’s Kansas City National SecurityCampus, which is operated and managed by Honeywell Federal Manufacturing & Technologies, LLC under contract number DE-NA0002839.1295References[1]K. Johnson, J. Blough, and A. Barnard, “Frequency Response Identification of Additively Manufactured Parts for Defect Identification,” in Annual International Solid Freeform Fabrication Symposium, August 2018, Austin, TX [Online].

namic Resolution Control in a Laser Projection based Stereolithography SystemYayue Pan*, Chintan DagliDepartment of Mechanical and Industrial Engineering, University of Illinois at Chicago, Chicago, IL 60607Corresponding author: yayuepan@uic.edu, (312)996-8777AbstractIn a typical Additive Manufacturing system, it is critical to make a trade-off between the resolution and build area for applications in which varied dimensional sizes, feature sizes, and accuracies are desired. The lack of the capability in adjusting resolution dynamically during building processes limits the use of AM in fabricating complex structures with big layer areas and small features. In this paper, a novel AM system with dynamic resolution control by integrating a laser projection in vat photopolymerization process is presented. Theoretical models and parameter characterizations are presented for the developed AM system.

Theoretical models and parameter characterizations are presented for the developed AM system. Accordingly, the process planning and mask image planning approaches for fabricating models with varied dimensional sizes and feature sizes have been developed. Multiple test cases based on various types of structures have been performed.Keywords: Additive Manufacturing, vat photopolymerization, Stereolithography, Resolution Control, Build Size1.

Multiple test cases based on various types of structures have been performed.Keywords: Additive Manufacturing, vat photopolymerization, Stereolithography, Resolution Control, Build Size1. Introduction1.1 Backgrounds and motivationsAdditive Manufacturing (AM), often referred to as Solid Freeform Fabrication (SFF) or 3D Printing or Rapid Prototyping, is a class of technologies that can fabricate parts with almost any freeform geometries directly from three-dimensional computer aided models by accumulating material together, usually in a layer by layer manner.In past decades, intensive research attempts have been made to improve the performance of AM systems, in terms of build speed [1, 2], surface quality [3-5], material property[6-8], process reliability [7], etc.

This technology finds its applications as a rapid prototyping technology in a wide variety of fields such as architecture, construction, industrial design, automotive, aerospace, military, engineering, medical, biotech, fashion, footwear, jewelry and education [9-11]. However, many limitations still exist in AM and hinder its application in fabricating end-use products, including size and resolution limitations, imperfections, limited choices of materials, low through-put, etc. [12]. In this research, we focus on working to overcome one of its primarydrawbacks: size and resolution limitations, in a Stereolithography related AM system which has entered the mainstream of 3D printing industry.Stereolithography (SL) related AM technology can achieve a wide range of resolutions. As of today, various commercial machines have successfully entered the market, typically offering resolutions in the range of 20µm ~200µm and build envelope diagonal size of 30 ~ 500 mm.

As of today, various commercial machines have successfully entered the market, typically offering resolutions in the range of 20µm ~200µm and build envelope diagonal size of 30 ~ 500 mm. There are two types of Stereolithography related AM technologies: laser scanning based, and projection based technologies. In the early stage, laser beam is used in SL system to cure liquid resin path by path to form one layer. With the advancement in Micro-Electro-Mechanical Systems (MEMS), projection based SL process becomes a more popular SL related AM technology these years due to their dynamic mask generation capability. In a projection SL system, a digital micromirror device (DMD) is usually used to pattern the light. A DMD is a Micro-Electro-Mechanical Systems device that enables one to simultaneously control small mirrors that turn on or off a pixel at over 5 KHz.

A DMD is a Micro-Electro-Mechanical Systems device that enables one to simultaneously control small mirrors that turn on or off a pixel at over 5 KHz. Using this technology, a light projection device can project a dynamically defined mask image onto a resin surface to selectively cure liquid resin into layers of the object. Consequently, the related AM process, Mask Image Projection Stereolithography (MIP-SL), can be much faster than the laser based SL process by simultaneously forming the shape of a whole layer, and can achieve a much higher resolution by having a big number of micro-mirrors and focusing each pixel in a small area. Many research groups demonstrated micro-manufacturing capability with micron or even sub-micron scale resolution in MIP-SL systems [3, 5, 13-15].Despite the advancement of SL related AM technologies, conflict between the build size and resolution is a historically inherent manufacturing dilemma.

In order to achieve higher resolution, a smaller laser beam spot or a smaller pixel size is desired, in turn, a smaller build size will be created. Due to the fundamental energy distribution mechanism in SL related AM technologies, build size and resolution have always been two conflicting goals. Recently a method to solve this dilemma by projecting multiple images in different areas to make one big layer is proposed by Emami et. al [16]. However this approach will lengthen the build time greatly and make the process much more complicated.The thrust of this research is to contribute to the advancement of AM by addressing such a dilemma of the resolution and build size, without affecting other manufacturing performance factors like build speed and process reliability.

To achieve this goal, a continuous resolution control approach is investigated by using a laser projection technology.1.2 Novelty and ContributionsA schematic diagram of the conventional MIP-SL process is represented in Fig. 1a. A lamp is usually used as a light source and a DMD chip is used to pattern the light. Optical lens is used to focus the patterned light onto a liquid resin surface. In a system, the resolution is determined by the lens' focal length and fixed. To cure a layer with different area, different number of pixelscould be illuminated. As shown in Fig. 1b, with the conventional approach, the curing area of a layer could be adjusted along a line segment by adjusting the number of illuminated pixels. For example, assume there are 1024 x 768 pixels in total and the pixel size is r2, the curing area that could be adjusted by changing the number of illuminated pixels and the control freedom is the green line segment.

In order to enlarge the build volume, the pixel size r2 should be increased which in turn will reduce the achievable resolution, causing the dilemma between resolution and build size.Light SourceLensResin VatBuild Physical ModelLiquid ResinZ-stagePlatformDigital Micromirror Devicer1Our ApproachBuilding AreaNumber of PixelNA1024 x 768r3r2Conventional Approach(a) (b)Figure 1. (a) Schematic illustration of conventional MIP-SL systems (b) Pixel number- Curing area relation in conventional approach and our approachTo address this problem, we propose a dynamic and continuous resolution control approach for MIP-SL systems. With our approach, the pixel size can be adjusted dynamically within a certain range from r1 to r3. Consequently, the control freedom is a 2-D area, instead of a 1-D line segment, as the light blue area showed in Fig 1b.This novel dynamic resolution control approach is presented in the following sections.

Section 2 presents an overview of a novel Laser-Projection based Stereolithography (SL) System with the resolution control approach along with its flowchart explaining the working of the system. A special Mask Image Processing Algorithm for dynamic resolution control is presented in Section 3, followed with a parametric dependence quantification process investigated in Section 4. Multiple test cases are demonstrated and discussed in Section 5. It is shown that parts with varying layer sizes and feature sizes could be fabricated with desired resolutions, by using this approach. Finally, conclusions are made in Section 6.2. Overview of the Laser Projection based SL SystemInstead of a lamp, a laser is used as a projection light source in our system. The laser projector has unique properties including focus free operation, capability to produce dynamic mask image irrespective of any surface (flat or curved) as shown in Fig. 2.

The laser projector has unique properties including focus free operation, capability to produce dynamic mask image irrespective of any surface (flat or curved) as shown in Fig. 2. And last but not the least is itscompact size that makes it easy to be adopted and integrated into the additive manufacturing system. These unique characteristics of laser projector eventually provide a distinct capability to change the projection image size and resolution easily.The laser projector is composed of three LEDs (blue, red and green) and a MEMS scanner. The projected image is created by modulating the three lasers synchronously with the position of the scanned beam [17]. The projected beam directly leaves the MEMS scanner and creates a sharp image irrespective of any surface (flat or curved) it is shone upon [17].(a)(b)(a) (b)Figure 2.

(a) A Laser Projection based SL system with dynamic resolution control (b) Flowchart of the working of Laser Projection based SL systemThe Laser Projection based SL system design is shown in Fig. 3a above and the flowchart in Fig. 3b explains the working of the entire system. Similar to the conventional approach, we start from a CAD model. However, instead of slicing the STL file with a certain predefined resolution setting, we first analyze feature sizes of the CAD model, and then set a resolution for each layer. After that, mask images will be generated by slicing each layer with the desired resolution setting. These mask images are then transferred to the laser projector for layer by layer fabrication. The laser projector then moves to a corresponding position to cure that layer with thedesired resolution and simultaneously the pixel size changes dynamically according to the position of laser projector.

Therefore, in this dynamic resolution control manufacturing process the above two steps highlighted in blue are the most important. Correspondingly, two fundamental research questions need to be answered:1. How to determine a proper set of resolutions for each layer to achieve best part quality.2. How to move projector to the proper position to gain the required resolution.Question 1 will be answered in Section 3 by exploring an image processing algorithm to filter features with different resolution requirements. Question 2 will be investigated in Section 4 by quantifying the parametric dependence of resolution and build size on manufacturing process parameters.3. Image Processing AlgorithmIt is therefore very essential to answer the first question about how to determine proper resolution for each layer so as to obtain best quality part and therefore we proposed an image processing algorithm to answer this question.

For instance, a solid freeform fabrication design model is sliced into a set of layers and it might have micro-scale and a macro-scale features in a single layer. If this layer is fabricated with a proper resolution corresponding to the macro-scale feature, the micro feature could not be fabricated out due to the insufficient resolution. However, if a high resolution setting is used to give the required accuracy for fabricating micro features, the build size may not be big enough to fabricate out the whole layer. Therefore, for each sliced layer, a proper resolution or a proper set of resolutions need to be determined to achieve the desired accuracy and build size for that layer without sacrificing the other manufacturing performances like build speed.In this study, a single Layered Depth Image (LDI) algorithm is utilized to determine the proper set of resolutions for each layer.

LDI is generally used for 2-dimensional images which can be extended further to Layered Depth-Normal Images commonly known as LDNI [18]. In the following sections, for easy understanding, we use "micro feature" or "micro image" to denote an area which has to use a high resolution R1 to fabricate, and "macro feature" or "macro image" to denote an area which could use a low resolution R2 to fabricate.The following example Fig. 4 illustrates a simple sliced binary image for a single layer that contains micro feature in a large area. We segmented the micro features from the macro part so as to build these small features with higher resolution in order to highlight its intricate details. Let V comprises of the entire image having i pixels along X-direction and j-pixels along Y- direction. Originally, the image is sliced by using the lowest resolution setting in our system, which is 81 microns per pixel denoted as R81.

Originally, the image is sliced by using the lowest resolution setting in our system, which is 81 microns per pixel denoted as R81. This can be represented as (Xi, Yj) ∈ V for a single layer that can be extended to the 3rd dimension identifying the Kth sliced layer. Thus the overall equation can be represented as (Xi, Yj) ∈ VLayer_K. In order to filter the micro feature which cannotbe fabricated by using the lowest resolution R81, rays are transmitted along each row pixel first and then each column pixel. We consider each row as a separate image and run the loop. When the ray identifies a pixel that has its neighboring pixel of different intensity it shall be marked as the boundary of the object.In the following example, a sectional part is highlighted to show the working of the algorithm. The system identifies the object pixel and mark the boundary pixel at the respective position as Xi,n, where n is the change in the pixel intensity ranges from 0 to infinity.

The system identifies the object pixel and mark the boundary pixel at the respective position as Xi,n, where n is the change in the pixel intensity ranges from 0 to infinity. The ray starts with n value as 0 and changes to 1 immediately when it finds the 1st boundary pixel and then proceeds with an increment when it encounters discontinuity. Thus, the next boundary will have Xi+p,n+1, where p is any arbitrary value of X position. The distance between any two boundaries of an object can be generalized as d|Xi+p,n_even- Xi,(n_even)-1|, except when n is 0.

The distance between any two boundaries of an object can be generalized as d|Xi+p,n_even- Xi,(n_even)-1|, except when n is 0. We set a threshold value T for the pixel distance which indicates that if the distance between two boundary object pixel exceeds the threshold it will be counted as meso-scale or macro-scale feature, whereas if the distance between the boundary object pixel is less than the threshold it will be counted as the micro-scale feature which cannot be fabricated by using resolution R81 and the object pixel will be marked as 0.X0,0XYX5,1X15,2X29,3X35,4X49,5X59,6(a)(b)(d) (e)(c)Figure 4. Example Model (a) STL Model; (b) Transfer Rays along Rows and Columns of a Binary Sliced Layered Image; (c) Sectional View highlighting object and boundary detection; (d) Macro Image; (e) Micro ImageFor example, in the sectional image in Fig. 4, the ray transmitted along the row starts with n as 0 and then ends with n as 6, thus the row has 6 boundaries and 3 objects.

4, the ray transmitted along the row starts with n as 0 and then ends with n as 6, thus the row has 6 boundaries and 3 objects. It is identified that the d|X15,2 – X5,1| and d|X59,6 – X49,5| is greater than the threshold T, thus identified as macro image. On the contrary, the other object is micro image since the d| X35,4 – X29,3| is less than T. Afteridentifying the micro-scale features, it will set all micro pixel object area to zero, we run the loop for all the rows and then a new image is generated. For better efficiency we then run this loop for all the columns on the new image and again the identified micro objects are set zero forming a modified image. The modified image thus obtained is the macro image and subtracting this modified image from the initial image will give out a micro image. Then the macro image is used for fabricating the related part by using R81.

Then the macro image is used for fabricating the related part by using R81. The micro features separated from the original sliced image, a higher resolution R81- (denoting a resolution of 81- microns per pixel) is used to regenerate an image for the micro features. Then the regenerated image is used as the input image and is processed the same way to determine a proper resolution for it. Following algorithm provides a throughout understanding of the working of image processing approach. R# means a resolution of # microns per pixel.Algorithm1: INPUT A binary image sliced by using R#, V of Kth layer. OUTPUT 1. Macro Image 2. Micro Image STEPS1. Transfer rays from Yj till Yj+w, where j=0 w and w = number of pixels along Y-axis 2. For each row find boundary pixel, Xi,n, where i=0 v and v= number of pixels along X- axis and n= change in pixel intensity3. If d|Xi+p,n_even- Xi,(n_even)-1|≥ T, macro image else micro-image 4.

For each row find boundary pixel, Xi,n, where i=0 v and v= number of pixels along X- axis and n= change in pixel intensity3. If d|Xi+p,n_even- Xi,(n_even)-1|≥ T, macro image else micro-image 4. If d|Xi+p,n_even- Xi,(n_even)-1|<T, set d|Xi+p,n_even- Xi,(n_even)-1|=0 5. New image as V_row 6. Repeat above steps for column rays on V_row 7. Store the modified image V_mac 8. V_mic =V - V_macRESULT V_mac as macro image which will be used to fabricate a part of that layer using resolution R# V_mic as micro image which the resolution needs to be updated.4. Parametric Dependence QuantificationAfter obtaining the desired resolutions for each layer using Image Processing Algorithm, the next thing is to move the laser projector to the corresponding position to fabricate that part of that layer with its desired resolution. This is quantified by the dependence of pixel size and build size on the focal length. By moving the laser projector up and down using Z-stage 2 as shown in Fig.

This is quantified by the dependence of pixel size and build size on the focal length. By moving the laser projector up and down using Z-stage 2 as shown in Fig. 3a, the digital mask images exposed on the liquid resin would have changing resolution and build size. More specifically if the laser projector is moved away from the platform the pixel size increases which means lower resolution but bigger build area. For example, the micro scalestructures could be segmented and fabricated with a small focal length and hence higher resolution, while the meso-scale structures in the model could be fabricated with a bigger focal length and hence larger curing area.

The following section shows the relationship between pixel size and build size on the focal length.4.1 Quantifying Parametric Dependence of: Build Area & Resolution on Focal LengthDue to the laser property of producing focus free masks irrespective of any focal distance, it provides a wide variety of resolution depending upon its distance from the bottom surface of resin vat. Thus, a process planning approach is essential for determining the position of the projector to obtain the required resolution for each layer. Both the resolution and building area are related to the focal length and there exist a linear co-relationship which has been highlighted in the following Fig. 5.

Both the resolution and building area are related to the focal length and there exist a linear co-relationship which has been highlighted in the following Fig. 5. The graph highlighted in red is related to the focal length and the resolution (y-axis on the right termed as 𝑦2) it produces whereas the graphs in blue (dashed line) and orange (round dotted line) are associated to the focal length and building sizes (y axis on the left termed as 𝑦1) along X and Y-axis respectively.

The linear relationship between focal length x and building area y could be formulated as the following:𝑦1 = 0.8 ∗ 𝑥 + 14.1333 along X-axis𝑦1 = 0.8 ∗ 𝑥 + 14.1333 along Y-axiswhereas, Focal Length x (mm) and Planar Resolution y (micron/pixel) could be described by:𝑦2 = 0.9906 ∗ 𝑥 + 16.6597From the above equations, it's obvious that for producing parts with high resolution the focal distance between the laser projector and resin vat should be small and conversely for the development of parts with large area its focal distance will be large. Given a desired resolution, the projector could be moved to the corresponding position. Also, the fabrication of one layer with large area but small features, the mask image could be segmented into multiple images for multiple curing with different projector positions. As shown in Fig.

Also, the fabrication of one layer with large area but small features, the mask image could be segmented into multiple images for multiple curing with different projector positions. As shown in Fig. 5, the projector provides a good flexibility with a building area of 73.0 mm to 31.0 mm along X-axis and 41.32 mm to 17.54 mm along Y-axis subject to the focal length of 70.0 mm to 20.0 mm respectively. Correspondingly, its resolution varies from 86.08 micron/pixel to 36.55 micron/pixel.Focal Length (mm)Building Size (mm)Resolution (micron/pixel)Focal Length vs Resolution (y=0.9906*x + 16.6597)Focal Length vs Building Size along X-axis (y=0.84*x + 14.1333)Focal Length vs Building Size along Y-axis (y=0.4756*x + 7.9913)010203040506070809001020304050607080901001015202530354045505560657075Focal Length vs Building Area vs ResolutionFigure 5.

Graphical Relationship between Focal Length, Building Area and Planar Resolution4.2 Validating Beer Lambert EquationTo measure cure depth (Cd) experiments were conducted for further analysis of the vertical resolution. The laser projector produced a mask at different focal lengths with each having a set of exposure time of 200 s, 300 s, 400 s, 500 s and 600 s. A commercial resin R5 red from Envisiontec containing 0.1-5% photoinitiator was utilized for the following set of experiments [19]. Other resin parameters such as penetration depth Dp and critical exposure Ec can be obtained from the Beer-Lambert formula provided in equation below [20].𝐶𝑑 = 𝐷𝑝 ∗ ln (𝐸𝑚𝑎𝑥 𝐸𝑐) ……….. (1)Also, the values of Dp and Ec can be verified from the graph below. The graph shows the relation between cure depth and exposure energy per unit area on the resin surface. The results indicate the cure depth is linearly proportional to the natural logarithm of the energy per unit area.

The results indicate the cure depth is linearly proportional to the natural logarithm of the energy per unit area. Theoretically, the slope of the line is its penetration depth Dp and Ec is the energy when cure depth is 0, but this does not indicate the gel point of the resin [21]. It is observed from the graph that the value of Dp is 0.266 mm whereas Ec is 23.46 mJ/cm2.00.10.20.30.40.50.60.70.80.92.7533.253.53.7544.254.54.7555.255.55.7566.256.5Cure Depth, Cd (mm)Natural Logarithm of Exposure Energy, Emax(mJ/cm2) Cure Depth vs Exposure EnergyFigure 6. Graphical Relationship between Exposure Energy and Cure Depth4.2.1 Dependence of Cure Depth on Focal LengthOn further extending Beer Lambert equation the relationship between the cure depth and focallength has been established using equation 1, as represented below:𝐶𝑑 = 𝐷𝑝 ∗ ln (𝐸𝑚𝑎𝑥 𝐸𝑐)𝐶𝑑 = 𝐷𝑝 ∗ ln (𝑃𝑚𝑎𝑥 ∗𝑇 𝐴∗𝐸𝑐)……….

(2)where, Pmax = Laser Power, T= time, A= cured area.Since all the variables in equation 2 are constant for a specific time and also the cured area isdirectly proportional to the focal length d, the above equation can be represented as follows:𝐶𝑑 = 𝐷𝑝 ∗ 𝑙𝑛 (𝐾𝑐𝑜𝑛𝑠𝑡𝑎𝑛𝑡 𝑑)𝐶𝑑 = 𝐷𝑝 ∗ ln(𝐾𝑐𝑜𝑛𝑠𝑡𝑎𝑛𝑡) − 𝐷𝑝 ∗ ln(𝑑)𝐶𝑑 = 𝐾′𝑐𝑜𝑛𝑠𝑡𝑎𝑛𝑡 − 𝐷𝑝 ∗ ln(𝑑) ……… (3)Equation 3 can be validated with the graph below that justifies the relation between cure depth and focal length. These linear equations in Fig. 7 are obtained when the laser projector produce slice image on the build surface for a specified time under different focal length.Time 200 secTime 300 secTime 400 secTime 500 secTime 600 sec𝑦=−0. 66∗𝑥+1.565𝑦=−0. 66∗𝑥+1.515𝑦=−0. 66∗𝑥+1.46𝑦=−0. 66∗𝑥+1.395𝑦=−0. 66∗𝑥+1.340.10.20.30.40.50.60.70.80.912.72.93.13.33.53.73.94.14.34.5Cure Depth, Cd (mm)Natural Logarithm of Focal Length (mm)Cure Depth vs Focal LengthFigure 7. Graphical Relationship between Focal Length and Cure Depth5.

Graphical Relationship between Focal Length and Cure Depth5. Experimental Results and DiscussionTo validate the efficacy of the approach a testbed was developed and models with multiple resolutions, varied dimensional sizes were fabricated. These 3D STL models were initially sliced layer by layer into 2D images and projected through a portable laser projector. The unique feature of these experiments was providing dynamic motion to the laser projector during building process. This resulted in providing flexibility to fabricate parts with changing resolution in a single building task.5.1 Large Area Fabrication CapabilityThe manufacturing capability of fabricating parts with large layer areas was tested and verified by fabricating a spur gear. The CAD model and fabricated result are shown in Fig. 8. After analyzing this model using our image processing algorithm it was found that the part can be fabricated with one resolution.

The CAD model and fabricated result are shown in Fig. 8. After analyzing this model using our image processing algorithm it was found that the part can be fabricated with one resolution. A resolution of 81 micron/pixel was adopted which is good enough to produce all details of this gear model. Based on the resolution the laser projector’s focal length was kept 65 mm so as to obtain the desired resolution.(a)(b)(c)Figure 8. Gear:- (a) Isometric view of STL Model; (b) Isometric view of Built Part; (c) Top View of Built PartThis fabricated model constituted of 50 layers in total, including 7 base layers with a layer thickness of 90 microns. Initial exposure time for the base layer was 600 seconds per layer whereas for the subsequent layer it was 70 seconds each.5.2 Small Area Fabrication CapabilityThe system showed appropriate results for large area fabrication. The same experimental setup was adopted to test the limits of resolution and fabricated parts with the small area.

The same experimental setup was adopted to test the limits of resolution and fabricated parts with the small area. Also, to further endorse the system capability, minute details were added in the part. The CAD model as demonstrated in Fig. 9a is a hollow ball with nine circular cuts on its circumference. Four symmetric bigger circular cuts along the center whereas slightly above it includes four circular cuts of slightly smaller diameter. Lastly, the top portion is drilled to form a small circle. The image processing analysis showed that a single resolution, 30 micron/pixel works for fabricating the whole part with desired accuracy while a shortest build time. Thus, the focal length was fixed at 13.5 mm to give a 30 um/pixel resolution for building all layers.0.2mm0.5mm(a)(b)(c)Figure 9.

Thus, the focal length was fixed at 13.5 mm to give a 30 um/pixel resolution for building all layers.0.2mm0.5mm(a)(b)(c)Figure 9. Hollow Ball; (a) Isometric view of STL Model; (b) Microscopic imaging of Isometric view of Built Part; (c) Microscopic imaging of Top view of Built PartThe fabricated part, thus formed was 2.8 mm in diameter and 2.65 mm in height with bigger and smaller circular cuts of diameter 800 μm and 600 μm respectively. Also, the top circular portion had a diameter of 70 μm. The Fig. 9b and 9c are microscopic images of the isometric and top view of the part respectively. The entire structure was comprised of 63 layers including 7 base layers. Each layer corresponds to a thickness of 43 microns.

The entire structure was comprised of 63 layers including 7 base layers. Each layer corresponds to a thickness of 43 microns. Due to the reduced focal length the initial exposure time and the exposure time for both base and subsequent layers were kept as low as possible mainly 200 seconds and 30 seconds respectively.5.3 Multi-Scale Resolution Fabrication CapabilityThe system successfully fabricated parts with micro and macro scale features separately. The following test case is conducted to verify the manufacturing capability of building multi scale features in a single building task. For this purpose a gear set with concentric center was designed as shown in Fig. 10a. With our dynamic resolution control method, after applying the image processing algorithm, we identified that for the bottom gear, a pixel size of 81 micron is good enough to give the desired accuracy and build size.

Whereas for the top gear, since its feature is much smaller, a smaller pixel size, 41 micron is desired to fabricate the features precisely. Aftersetting up different resolutions for layers, we could assign the corresponding projector position. According to the parametric dependence developed in this research the focal length was obtained as 65 mm and 25 mm for the bottom and top gear respectively.(a)(b)(c)Figure 10. Multi-scale Gear; (a) Isometric view of STL Model; (b) Isometric view of Built Part; (c) Top view of Built PartThe fabricated part constituted 39 layers including 7 base layers with a layer thickness of 90 micron. The base layer were cured for 600 seconds were the subsequent layers were cured for 70 seconds.

The base layer were cured for 600 seconds were the subsequent layers were cured for 70 seconds. The bottom gear diameter comprised of 30.0 mm whereas the top gear diameter comprised of 5.00 mm with an overall depth of 3.8 mm.5.4 Continuous Resolution Control Fabrication CapabilityThe first two test cases in Section 5.1 and 5.2 demonstrate the capability of our approach on fabricating parts with different resolutions. Section 5.3 verifies the effectiveness and efficiency of our approach on fabricating a part with two different resolution settings in one build.Furthermore, another test case is conducted to demonstrate the system capability in continuously changing the resolution for each layer. The STL file shown below in Fig. 11a consists of a triangular structure with a circular hole drilled through its center. Also, the two sides of the triangle are drafted by an angle of 55 degree.

The STL file shown below in Fig. 11a consists of a triangular structure with a circular hole drilled through its center. Also, the two sides of the triangle are drafted by an angle of 55 degree. Since the layer by layer sectional area of the part reduced continuously, the projector’s focal length was also changed from 45.0 mm to 25.0 mm ultimately improving the resolution by 0.34 micron/pixel for each layer. The projector’s focal distance was controlled by synchronizing its motion with the platform.The model comprised of 59 layers in total with each layer thickness corresponding to 76 micron. The base layer of the model was fabricated with a curing time of 300 seconds whereas the subsequent layers were cured for 30 seconds each. The isometric view along with part dimension is shown below in Fig. 11b and 11c. The side of the triangle is measured to be 18.0 mm whereas the circular hole has a diameter of 5.0 mm.

The isometric view along with part dimension is shown below in Fig. 11b and 11c. The side of the triangle is measured to be 18.0 mm whereas the circular hole has a diameter of 5.0 mm. The top tilted edge is 9mm long while the other two edges that correspond to the part thickness are measured to be 4.5 mm.(a)(b)(c)Figure 11. Titled Edges Triangular Surface; (a) Isometric View of STL Model; (b) Isometric View of Built Part with Side Measurement; (c) Isometric View of Built Part with Edge Measurement5.5 Building Performance StatisticsThe system demonstrated a unique characteristic of dynamic and continuously changing resolution. Also, the laser projector being so compact in size can be conveniently moved along the z stage without considering the factor of the getting a distorted image due to its focus free operation. To make a more straightforward comparison with conventional approaches, the building time for base layers is excluded in the part build time.

To make a more straightforward comparison with conventional approaches, the building time for base layers is excluded in the part build time. The table below shows the capability of this approach on fabricating parts with different layer areas and feature sizes efficiently through changing resolutions dynamically and even continuously for each layer.Table 1. Statistics of all test casesModel 1 (Large Area)Model 2 (Small Area)Model 3 (Multi-scale)Model 4 (Continuous resolution)X-Y Dimension (mm)Smallest feature (mm) Depth (mm) Total number of Layers35.0 x 35.0~5 4.5502.8 x 2.80.07 2.656330.0 x 30.0 (big gear) 5.0 x 5.0 (small gear) 1 3.53918.0 x 16.00.2 4.559Part Build Time (min)50204530Layer Thickness (microns)904390766. ConclusionA novel mask image projection based stereolithography (MIP-SL) system has been developed to achieve dynamic resolution control.

ConclusionA novel mask image projection based stereolithography (MIP-SL) system has been developed to achieve dynamic resolution control. Compared to conventional MIP-SL systems, the light source in our design includes a compact laser projector that was mounted on a z-stage, which resulted in changing focal length and in turn projection area and resolution. To facilitate the mask image planning problem in dynamic resolution control approach, an image segmentation algorithm has been developed to separate micro features from a sliced image. Parametric dependence of resolution and build size on focal length have been calibrated and modeled. Theeffectiveness and efficiency of the system has been verified with multiple test cases with various surface areas, feature sizes and structures.References1. Ha, Y.M.,etal.,“Mass production of 3-D microstructures using projectionmicrostereolithography” Journal of mechanical science and technology, 2008. 22(3): p. 514-521.2.

M FABRICATION AND CHARACTERIZATION OF HIGH PURITY NIOBIUM FOR SUPERCONDUCTOR APPLICATIONSC.A. Terrazas1, S.M. Gaytan1, J. Mireles1, P. Frigola2, D. Espalin, R.B. Wicker11. W.M. Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, Tx, 79968 2. RadiaBeam Technologies, Santa Monica, CA, 90404AbstractSuperconducting radio frequency (SRF) cavities are used to accelerate charged particles to near the speed of light for elemental studies. Currently, SRF cavities are typically fabricated using different forming processes including deep-drawing and spinning to mechanically shape niobium into the desired geometry. This research presents the development of processing parameters for high purity niobium (powder size range of 25-125μm) using electron beam melting additive manufacturing technology. Fabrication parameters were improved to obtain dense parts in a time-efficient manner.

Fabrication parameters were improved to obtain dense parts in a time-efficient manner. A specific procedure was used to maintain powder purity, and powder chemistry was monitored at different stages of fabrication. In addition, a series of experiments were performed to obtain 99.9% dense parts and a maximum building height of ~85mm.IntroductionThe Electron Beam Melting (EBM) process is an additive manufacturing (AM) technology used for the rapid manufacturing of near net-shape metal components. There are many advantages of this process over other AM techniques. For example, since the process is performed in vacuum, it allows for the fabrication of high melting point metals and reactive materials such as nickel, cobalt and titanium alloys while eliminating the negative effects of impurities or oxidation.

Also, EBM is considered a hot process because the electron beam is utilized prior to fabrication of each layer to preheat the powder metal bed resulting in the fabrication of parts that have minimal to no residual stresses [1]. The EBM process has been used mostly for medical and aerospace applications, although continued research is widening the spectrum of materials that can be processed as well as their applications.The EBM process has been commercially available since 2000 by the Swedish company Arcam (Arcam AB, Sweden). This process is catalogued as a powder bed fusion AM technology and it uses a stream of electrons to sinter and melt selectively in a layer-by-layer fashion. Metal powder particles (30µm – 120µm in diameter) are used as the precursor material for fabrication. In the process, layers of metal powder with a thickness in the range from 50-120 μm are evenly spread by a raking mechanism.

In the process, layers of metal powder with a thickness in the range from 50-120 μm are evenly spread by a raking mechanism. Then the electron beam is used to preheat the powder bed followed by the selective melting in specific regions through the bed using higher energy on the beam. The electron beam is accurately and rapidly steered to scan specific geometries with the use of electromagnetic lenses or coils positioned in the gun column. After processing of one layer, the build table is lowered in the Z direction by a thickness of one layer and the process repeats until the part is finished. The process is carried out in vacuum levels that are maintained at ~10-4 Torr or below. The main components in the Arcam EBM process (Figure 1) are the electron gun (1), electromagnetic coils (2), powder hoppers or containers from which powder is gravity fed (3), the raking arm (4) and the build stage (6).

Normally, the process uses a buildFigure 1 – Schematic of the Arcam electron beam melting systemsubstrate to commence fabrication and provide anchoring of the initial layers of the part being produced (5).SRF CavitiesIn the field of particle physics, superconducting radio frequency (SRF) devices are used to generate electric fields that accelerate elemental particles such as electrons to study their interaction upon collision. SRF cavities are key components in particle accelerators since they transfer the energy to the traveling particles, and their purity is critical as it directly affects the accelerating field that can be achieved during operation. Assemblies of multiple SRF cavities are normally employed in a particle physics laboratory setup. In the setup, voltage is applied intermittently to the cavity assembly while particles are injected from a source at a rate that is in phase with the voltage.

In the setup, voltage is applied intermittently to the cavity assembly while particles are injected from a source at a rate that is in phase with the voltage. As the voltage oscillates through the assembly, each cavity experiences a change in polarity that provides the energy kick and accelerates the particles.High purity niobium is the material of choice for fabrication of SRF cavities since the element has the highest critical temperature of all pure substances (9.2K or -263.95ºC) and it becomes superconducting when cooled below this point. Subtractive (i.e. machining) manufacturing is rarely used in fabrication of SRF cavities because of niobium’s upscale price.Therefore, three forming methods (Figure 2) are currently used in the fabrication of cavities which are deep drawing, spinning and hydroforming [2]. Deep drawn cavities are formed by stamping a thin (~3mm) niobium sheet into a set of dies.

Deep drawn cavities are formed by stamping a thin (~3mm) niobium sheet into a set of dies. The spinning operation is performed similarly by starting with a niobium disc of similar thickness that is bolted from the center and rotated. Then an adjustable mandrel is moved to apply pressure and form the shape of the cavity. Both techniques mentioned before have the limitation that only half-cavities with uniform thickness can be produced. Post-processing steps to form the full cavity are required, which include the welding of the half-cavities from the equator; a process that introduces variability in thickness and possible contamination of the precursor purified niobium. The third process employed is hydroforming where a single niobium tube is attached in between a set of dies and hydraulic pressure is applied at the interior of the tube to form the cavity profile.

The third process employed is hydroforming where a single niobium tube is attached in between a set of dies and hydraulic pressure is applied at the interior of the tube to form the cavity profile. This process has the advantage of producing multi-cavity designs but further development is necessary to achieve uniform thickness of the resulting parts [2].Issues with Current SRF CavitiesDuring experimental operation of cavity assemblies, there are many factors that affect their performance. From the standpoint of the mechanical stability of the parts, the development of the so called Lorentz forces cause the mechanical deformation of cavities. These forces appear as an outward pressure at the equator and an inward pressure at the iris in the cavity profile eventually causing its deformation (Figure 3).

These forces appear as an outward pressure at the equator and an inward pressure at the iris in the cavity profile eventually causing its deformation (Figure 3). These local deformations cause the detuning in the frequency of the cavity assembly ultimately reducing the accelerating field thata)b)c)Figure 2 – Traditional fabrication methods for SRF cavities. a) deep-drawing, b) spinning and c) hydroforming.Figure 3 – Lorentz forces acting in the wall of an SRF cavity causing its deformation.can be achieved. Welding of support structures and rings is usually performed as corrective measures to prevent detuning; however, this is done at the expense of introducing more complexity and variability to the components.

The use of different joining techniques such as tungsten inert gas welding (TIG), metal inert gas welding (MIG), or electron beam welding (EBW) also increases the risk for contamination at the surface of cavities.Other issues with detrimental factors to the electronic properties of niobium cavities are field emission and thermal breakdown (quenching), both problems related to the presence of impurities and defects in the cavity wall. In field emission, electrons are emitted from sites with impurities and their trajectories bent by the magnetic field causes them to strike the walls of other cavities and failure occurs by localized heat. In the case of quenching, small surface defects such as roughness or interstitial impurities cause unstable power dissipation leading to breakdown of the accelerating field [3].EBM-AM of SRF CavitiesIn this work, EBM was explored as an alternative fabrication process given the limitations of current methods to produce SRF niobium cavities.

EBM provides the ability to fabricate nearly fully dense components with complex designs. This is an important advantage for the fabrication of SRF cavities as it can be used to manufacture designs that exhibit non- uniform thickness and/or with integrated stiffening support structures to resist the deforming forces and frequency detuning that develops during experimentation (Figure 4). The technology also offers the potential to produce cavities and other components in a monolithic form greatly reducing the post-processing by welding and hence maintaining the purity of the niobium once fabricated in the final shape.Experimental MethodsThe focus of this work was to optimize the processing parameters for EBM fabrication of dense niobium components using precursor powder with particle sizes in the range from 25 to 125µm.

Also, part of the work focused on the implementation of a procedure for handling the niobium powder to minimize the exposure to environmental impurities and humidity (describedFigure 4 – CAD rendering of the cavity profile with varying thickness and integrated support structures to the wall of a cavityin the powder handling section). The chemistry of niobium was traced from the wrought stock used, after its fabrication into wire, after the plasma atomization of the wire, and after EBM- fabrication.The approach followed to improve the density of EBM niobium implemented the use of three sets of parameter combinations. Each experiment consisted of the fabrication of nine separate blocks or coupons with dimensions of 10x10x10mm forming a grid or matrix (Figure 5). The processing parameters from the melt explored were beam speed (mm/s), average current (mA) and focus offset (mA) of the electron beam.

The processing parameters from the melt explored were beam speed (mm/s), average current (mA) and focus offset (mA) of the electron beam. Across the rows of the grid, the speed and average current of the beam were varied while the focus offset was varied column wise. A total of three experiments were performed in which the fabrication parameters that provided the highest density were used for the next iteration.After fabrication of the 9 coupons with different parameters, the density of the parts was measured using a weight balance (1) with a density measuring kit including a plastic stand (2),Figure 5 – Schematic of fabrication parameter space explored during experiments to improve the density of EBM niobium.one bracket (3), one weight basket (4), and a nalgene beaker (5) to hold distilled water, as depicted by Figure 6. Measurements were done according to ASTM Standard B311-13 by the volume displacement method [4].

Measurements were done according to ASTM Standard B311-13 by the volume displacement method [4]. Weight measurements were made by weighing the parts in air and then in distilled water. The apparent density was calculated with the formula Density = (A x E)/(A - F). In this formula, A was the mass of the test specimen in air, F was the mass of the test specimen in water and E was the density of water at the working temperature. The percent relative density (%RD) for parts was calculated by dividing the obtained density values over the density for wrought reactor grade niobium of 8.57g/cm3.Since the surface roughness obtained from as-fabricated EBM parts is not optimized for SRF applications, a final experiment was done to measure the density of machined EBM niobium using the same setup described before.

The edges or contours for one EBM-fabricated bar with dimensions of 12 x 12 x 85 mm were machined using a computer numerical control34521Figure 6 –Weight balance and density measurement kit used for EBM niobium parts.(CNC) machine. Then, the part was cut into five equal sections and density measurements were taken for the parts. The values obtained provided a better indication for the density of the niobium when fabricated by the EBM process. Reported values in this work were the average from three measurements for each part.Powder HandlingTo maintain the purity of niobium, handling of the powder material was performed using a glovebox with a controlled N2 gas atmosphere. The unit (Figure 7a) injected N2 gas during operation maintaining the humidity at or below 5%. Inside the glovebox, a mechanical sifter (Figure 7b) was designed, fabricated (Figure 7c) and used to speed up the process to sieve thepowder for recycling purposes.

Inside the glovebox, a mechanical sifter (Figure 7b) was designed, fabricated (Figure 7c) and used to speed up the process to sieve thepowder for recycling purposes. The sifting process was carried out in ~20 minutes and the powder was poured manually into the hoppers of the A2 system.Ventilation of the machine was also performed with N2 gas by retrofitting a tank into the solenoid valve of the A2 machine. Once the system reached ambient temperature conditions, the N2 gas was used to ventilate the chamber, delaying the adsorption of humidity into the niobium powder which could make it prone to an increase in interstitial impurities once fabricated by EBM.a)b)c)Figure 7 – a) Glovebox unit used for handling and sifting niobium powder, b) CAD design of sifter aider, c) fabricated sifter used inside glovebox.Monitoring of Material PurityThe purity of niobium was traced using chemical analysis of the wrought bar, wire, atomized powder, and EBM-fabricated niobium.

The chemical analysis on several specimens were sent to Element Materials Technology (Element, USA), where element tracing was performed by inductively coupled plasma mass spectrometry (ICP), to detect light elements such as hydrogen (H), oxygen (O) and nitrogen (N), and LECO combustion for carbon (C). The purity of EBM-fabricated niobium was also investigated with the use of X-ray diffraction (XRD).An RGA-100 residual gas analyzer (SRS, USA) was also installed to evaluate the quality of the vacuum and thus the presence of impurities at the interior of the chamber (Figure 8). The RGA permitted monitoring of several gas species during fabrication using EBM.

The RGA permitted monitoring of several gas species during fabrication using EBM. Use of the RGA is important since some of those gases, which include H, N, O, He, water vapor (H2O), and carbon monoxide (CO2) can become sources for the formation of interstitial impurities causing detriment of the electronic properties of niobium in SRF applications.Figure 8 – RGA unit installed on the Arcam A2 system.ResultsApparent and Percent Relative DensitiesThe bar charts depicted in Figures 9 through 12 show the results obtained for apparent and relative densities for the experiments performed. In experiment 1, the highest apparent and percent relative densities were those of part number 1 measured at 8.48g/cm3 (98.9 %RD).In experiment 2, the processing parameter space explored was near those of part number 1 in experiment 1. Improved density values were obtained for parts 4, 5, and 6 with %RD values above 99%.

Improved density values were obtained for parts 4, 5, and 6 with %RD values above 99%. The highest density was that of part number 5 with measured values of 8.53g/cm3 (99.5%RD).Figure 9 – Results for density measurements in experiment 1.Figure 10 – Results for density measurements in experiment 2.Figure 11 – Results for density measurements for experiment 3.Figure 12 – Measured apparent and percent relative densities for EBM part with machined edges.Figure 13 – Chemical analysis history of niobiumResults for experiment 3 show a further improvement in the density values. The densest part was number 7 with values of 8.548g/cm3 (99.74% RD). The fabrication parameters for this specimen are a focus offset of 31 mA, an average current of 18.5 mA, and a beam speed of 172 mm/s for the melt themes. Since the fabrication parameters for this part yielded the best density in parts, further fabrication of parts was carried out with these parameters.

Since the fabrication parameters for this part yielded the best density in parts, further fabrication of parts was carried out with these parameters. As previously mentioned, the surface roughness is not optimal for SRF applications. The last experiment carried out consisted on machining the contours of a square bar measuring ~85mm and sectioning it into five equal parts. Density values obtained for all five sections averaged 8.55g/cm3 and three nearly fully dense parts were measured at 99.9%RD (Figure 12).Chemical AnalysisFrom the results of the chemical analysis it was deduced that the EBM process did not cause contamination of the niobium. By contrast, the atomization process to produce the powder from the niobium wire feedstock showed changes to the chemistry of the material, mostly as an increase in the titanium (Ti) content.

By contrast, the atomization process to produce the powder from the niobium wire feedstock showed changes to the chemistry of the material, mostly as an increase in the titanium (Ti) content. Other elements that also showed levels above the standardFigure 14 - Typical RGA plot observed during fabrication with niobium.(ASTM B393) [5], such as nickel (Ni) and iron (Fe), and those prone to form interstitial impurities such as hydrogen (H) and oxygen (O) (Figure 13).Residual Gas AnalysisThe monitoring of different gas species during fabrication was done using the RGA. Theplot of the partial pressures versus time shows they were maintained below the required limit ~10-5 Torr (Figure 14).XRD AnalysisThe XRD patterns obtained (Figure 15) for powder (a) and the fabricated EBM samples (b) also show a clean niobium spectrum with the preferred orientation planes.

A bcc structurea)b)Figure 15 – XRD spectra for a) powder and b) EBM niobium specimenswith lattice parameter of a= 3.3Å is representative for both the powder and the fabricated specimen. The fabricated specimen analyzed belongs to a plane parallel to the fabrication direction (z-axis) and XRD analysis shows a preferential directionality in the (110) orientation.ConclusionsIn this work, a design of experiments was implemented to improve the density of niobium parts fabricated by the EBM process. Three parameters (beam speed, beam current and focus offset) were varied in each experiment in an iterative manner until the percent relative density, was improved to values over 99.9%. The work also reported on the monitoring of the purity of niobium at different stages in the production of niobium which are: a) niobium ingot, b) wire from stock material, c) powder state after plasma atomization, and d) after fabrication using the EBM process.

Evidence that the material maintained the purity required by SRF applications after processing by EBM was obtained by the use of the RGA and the XRD analysis. Also, the use of chemical analysis for light elements such as H, N and O, and by LECO combustion totrace carbon (C) evidenced the purity of the precursor powder material before and after use in the EBM process.References[1] Larsson et al., 2003. Rapid Manufacturing with Electron Beam Melting (EBM) – Amanufacturing revolution? Proceedings of SFF 2003.[2] Padamsee, H., Knobloch, J., Hays, T., 1998. RF Superconductivity for Accelerators 2nd ed.Wiley. New York, NY. Print.[3] Palmieri, V. 1997. Fabrication of niobium cavities. Proceedings of SRF Workshop 1980.Karlsruhe, Germany.[4] ASTM International. ASTM B311-13. Standard test method for density of powdermetallurgy (PM) materials containing less than two percent porosity. ASTM International. West Conshohocken, PA.[5] ASTM International. ASTM B393.

Standard test method for density of powdermetallurgy (PM) materials containing less than two percent porosity. ASTM International. West Conshohocken, PA.[5] ASTM International. ASTM B393. B10 Committee, 2009. Specification for Niobium andNiobium Alloy Strip, Sheet, and Plate. West Conshohocken, PA..

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperTHE ECONOMICS OF BIG AREA ADDITIVE MANUFACTURINGB. K. Post*, R. F. Lind*, P. D. Lloyd*, V. Kunc**, J. M. Linhal** and L. J. Love*Manufacturing Systems Research Group, Oak Ridge National Laboratory, Knoxville, TN 378311 **Deposition Science and Technology Group, Oak Ridge National Laboratory, Knoxville, TN 37831AbstractCase studies on the economics of Additive Manufacturing (AM) suggest that processing time is the dominant cost in manufacturing. Most additive processes have similar performance metrics: small part sizes, low production rates and expensive feedstocks. Big Area Additive Manufacturing is based on transitioning polymer extrusion technology from a wire to a pellet feedstock.

Big Area Additive Manufacturing is based on transitioning polymer extrusion technology from a wire to a pellet feedstock. Utilizing pellets significantly increases deposition speed and lowers material cost by utilizing low cost injection molding feedstock. The use of carbon fiber reinforced polymers eliminates the need for a heated chamber, significantly reducing machine power requirements and size constraints. We hypothesize that the increase in productivity coupled with decrease in feedstock and energy costs will enable AM to become more competitive with conventional manufacturing processes for many applications.

As a test case, we compare the cost of using traditional fused deposition modeling (FDM) with BAAM for additively manufacturing composite tooling.IntroductionAccording to the Wohlers Report, sales of additive manufactured products and services was approximately $4.1B in 2014 and is expected to grow to $21B by 2020.(Wohlers 2015) This is an average growth rate of 39% per year. This corresponds to the recent history from 2010 to 2013 in which the industry experienced a growth rate between 24.1% and 35.2%. However, to put this in perspective, global manufacturing is approximately $12.8T per year suggesting that AM only represents 0.03% of worldwide manufacturing. Optimism suggests that there is room to grow whereas pessimism suggests AM is restricted to niche markets. An analysis of the current economics of AM suggests the latter.

Optimism suggests that there is room to grow whereas pessimism suggests AM is restricted to niche markets. An analysis of the current economics of AM suggests the latter. However, the authors hypothesize that it could be the former with disruptions in the technology.As a case in point, (Salmi 2012) provide a case study comparing high pressure die casting (HPDC) to selective laser sintering (SLS) to manufacture an aircraft component. To take advantage of the process, the team redesigned the component reducing the weight by 72% (e.g. the weight of the AM part was 3.6X lighter). An analysis of HPDC identified four primary cost elements: material, tooling, machine cost and labor for pre- and post-processing. In a similar manner, SLS had four primary cost elements: material, machine cost and labor for pre- and post-processing.

In a similar manner, SLS had four primary cost elements: material, machine cost and labor for pre- and post-processing. The fundamental difference between the two manufacturing approaches was the lack of tooling cost for SLS and the production rate (3.6 seconds per part for HPDC versus 13.5 hours for SLS). If cost is the primary metric for comparison, HPDC was $21.29+$21,000/N where $21,000 was tooling cost and N is part count whereas SLS had a fixed cost of $526.31 per part. Therefore, if the production volume is less than 41 parts, SLS is the more economical manufacturing process. At a volume of 41 parts, tooling accounts for 97% of the manufacturing cost in HPDC but decreases as volume increases. For SLS, machine cost per part (e.g. time) accounts for 90% of the total part cost and does not change with volume (see Table 1).

For SLS, machine cost per part (e.g. time) accounts for 90% of the total part cost and does not change with volume (see Table 1). Therefore, time, not material or labor, is the most expensive cost of additive manufacturing.Notice of Copyright This manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes.

The Department of Energy will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).11176Table 1: SLS Additive Cost BreakdownPreprocessing Material Processing Post processingCost $8.00 $25.81 $472.50 $20.00Percentage 1.5% 4.9% 89.8% 3.8%(Wohlers 2015) provided a similar case study conducted by Johnson Controls. The part in the investigation was approximately 250 x 120 x 50 mm and is currently manufactured via injection molding (IM) with a production volume of 8500 units per year. The part was redesigned for SLS with no requirements for finishing or post- processing. The unit cost for injection molding was $2.64 whereas it was $39.42 per unit using SLS. In this case, tooling accounted for 25% of the IM part cost whereas material and time accounted for 96% of the cost of the SLS part.

In this case, tooling accounted for 25% of the IM part cost whereas material and time accounted for 96% of the cost of the SLS part. Aggressive decreases in material cost and manufacturing time (order of magnitude) show that there is a pathway towards cost competitiveness between additive manufacturing and conventional manufacturing. The critical question is whether these disruptions are theoretically possible. Our hypothesis is yes if there are fundamental changes in both the technology (orders of magnitude increase in speed) and the business models (orders of magnitude drop in material cost).State of the ArtTable 1 provides a survey of typical build volumes and rates for commercial AM systems. In terms of build rate, extrusion technologies (extrusion, jetting, directed energy) are relatively straightforward to estimate since is it directly related to flow rate of the material. Estimate of production rates for powder bed and vat systems are slightly more complex.

Estimate of production rates for powder bed and vat systems are slightly more complex. The layer time is a function of both fusion and the preparation of the bed (raking and possibly preheating the material). Therefore the majority of industry quoted rates is measured in layer time rather than volumetric rate.

Therefore the majority of industry quoted rates is measured in layer time rather than volumetric rate. In order to provide a rough order of magnitude comparison between technologies, we bound the volumetric rate by assuming between 2% and 20% of the layer volume is fused in powder bed systems.Table 2: AM Volumes and RatesSystemTechnologyArcam Q20 Concept Laser Xline 2000R SLM Solution 280 HL Stratasys 900 mc ExOne mFlex Stratasys Connex 500 Optomec MR-7 3D Systems ProX 950 Fabrisonic SonicLayer 4000 Sheet laminationPowder bed fusion Powder bed fusion Powder bed fusion Material Extrusion Binder jetting Material jetting Direct energy Vat polymerizationBuild Volume (cm x cm x cm) 35 dia x 38 tall 80 x 40 x 50 28 x 28 x 35 91 x 61 x 91 40 x 25 x 25 50 x 40 x 20 30 x 30 x 30 150 x 75 x 55 100 x 60 x 60Build Rate (cm3/hr) 20 to 400 10 to 100 20 to 35 10 to 80 12 to 260 50 to 500 20 50 to 600 500The rates quoted in Table 1 only consider the manufacturing process.

Pre- and post-processing likewise impact total production rate. In the case of electron beam powder bed systems, the unmelted material is lightly sintered and must be mechanically removed, broken down and filtered prior to reuse. Laser powder bed systems do not preheat the powder to the level of the electron beam systems so powder separation is easier. However, the resulting increased temperature gradients result in higher residual stress requiring welding the parts to the start plate and heat treatment prior to part removal. Binder jet systems generate green parts that require sintering in a furnace prior to use. Material extrusion, vat polymerization and material jetting systems produce near net shaped parts but must have support material removed either mechanically or chemically prior to use. Sheet lamination processes are additive/subtractive systems that require machining processes to remove material in-1177situ to the process.

Sheet lamination processes are additive/subtractive systems that require machining processes to remove material in-1177situ to the process. The magnitude of post processing varies from process to process. However, in general, the additive process is the main cost element in additive manufacturing and are not currently competitive with traditional manufacturing processes.Big Area Additive ManufacturingConventional fused deposition modeling is based on melting and extruding a filament of thermoplastic feedstock. Prior work shows that the peak flow rate is limited by the rate at which the filament can be melted.(Monzon 2013) Big Area Additive Manufacturing is an extrusion process that uses injection molding material for the feedstock and a single screw extruder for melting and metering the flow rate (see Figure 1). (Holshouser, Newell et al.

(Holshouser, Newell et al. 2013) A gantry system, commercialized by Cincinnati Incorporated (see Figure 2), moves the extruder in x, y and z directions to build the part. The extruder is capable of delivering 100 lbs/hour of thermoplastic materials from pellet feedstock. The gantry system is capable of achieving 200 inch/sec peak velocities with 64.4 in/s2 accelerations and position accuracy in the 0.002”. As previously reported, the use of carbon fiber reinforcement in the thermoplastic resins increases the part strength and stiffness.(Tekinalp, Kunc et al. 2014) But, just as important, it increases thermal conductivity and reduces the coefficient of thermal expansion relaxing the need for a heated chamber to produce large parts (see Figure 3 and Figure 4). (Love, Kunc et al. 2014) The elimination of the oven significantly decreases the energy intensity (manufacturing energy required per kg of product).

(Love, Kunc et al. 2014) The elimination of the oven significantly decreases the energy intensity (manufacturing energy required per kg of product). As shown in Figure 5, conventional FDM systems with heated chambers have a 100 kW-hr/kg energy intensity. Desktop systems that have similar production rates have a 5.5 kW-hr/kg energy intensity suggesting that the oven accounts for 95% of the energy utilization in FDM production. BAAM further reduces the energy intensity to 1.1 kW-hr/kg by significantly increasing productivity from 1 ci/hr to 2500 ci/hr and manufactured at room temperature.1178Figure 1: BAAM ExtruderFigure 3: Section of a wind turbine mold1179Figure 2: Cincinnati BAAMFigure 4: Printed prototype houseFigure 5: Energy intensity of manufacturingTherefore, the combination of lower energy intensity, higher productivity and lower material cost suggests that there will be a significant production cost reduction with BAAM when compared to conventional AM systems.

As with conventional additive manufacturing, there are four primary costs associated with manufacturing a part: preprocessing (PR), material (M), processing (P) and post processing (PO). The combination of significantly higher production rates with low cost feedstock significantly changes the cost distribution of each of these elements. Preprocessing typically includes preparation of the file (slicing, support, orientation) and preparation of the machine (loading material and file). Material is generally slightly over the volume of the part and includes support material. Post processing includes removal of the part, support, recycling of any material and final finishing (machining, polishing…).Composite Tooling Case StudyTooling is the foundation of many manufacturing industries. Molds are used in the aerospace, automotive and appliance industry to shape metal, plastic and composite parts. As an example, the automotive industry generally spends $200M on tooling for each car model.

As an example, the automotive industry generally spends $200M on tooling for each car model. By 2018, tooling for the automotive industry is expected to exceed $15.2B per year. Tooling is generally very expensive ($10K’s to $100K’s per mold) and long lead (months to years) items. Therefore there is great potential for revolutionizing the tool and die industry through the use of additive manufacturing.The aerospace industry, due to weight concerns, has begun to migrate from aluminum to carbon fiber composites. The conventional approach for manufacturing composite structures is to layup the carbon fiber, core structure and resin on a mold. The molds and composite materials are processed inside an autoclave to rapidly cure the resins and remove any air. The requirements for the molds are elevated operating temperatures (from 200 C to 375 C), vacuum integrity (15 mbar over 60 minutes) and moderate pressures (100 psi) with tight dimensional tolerances (under 0.005” distortion).

A conventional composite tool will cost in excess of $100K (US) and take several months to manufacture. ORNL recently partnered with Boeing, Ford, Techmer and BSF to evaluate the feasibility and economics of using BAAM to produce composite tools. A variety of materials were evaluated including PEEK, PEEK, Ultem, Apollo, PPS, PPSF and PPSU. The final material of choice, based on performance and cost, was PPS loaded with 50% carbon fiber. The material cost was approximately $8/lb. The sample tool weighed approximately 65 lbs and took 2 hours to print. Preprocessing time (setting up 1180the machine and slicing the model) was approximately 2 hours. The mold was grown approximately 0.100” over sized and machined using a Thermwood router to the final size. It took approximately 2 hour to transfer the mold to the router and another 5 hours to machine. Assuming labor rates of $50/hr and machining rates of $150/hr, the total tool cost was approximately $1870 (see Table 3).

Assuming labor rates of $50/hr and machining rates of $150/hr, the total tool cost was approximately $1870 (see Table 3). In terms of manufacturing time, from beginning to end, it took approximately 2 days to fully manufacture the mold. In comparison to SLS in the BAAM provides a more balanced distribution of costs. AM machine time is equivalent to material cost and the cost of printing is close to the cost of finishing.Table 3: BAAM Tool Cost BreakdownPreprocessing Material Processing Post processing (labor) - Labor - machine TotalTime/material 2 hours 65 lbs 2 hours2 hours 5 hours 9 hoursCost $100 $520 $400$100 $750 $1870Percentage 5.3 27.8 21.4 45.5A similar analysis can be conducted for the same tool to be manufactured out of PPS on a Fortus 900 mc. First, material cost is approximately $650 for 93 cubic inches of material. The tool has 1725 cubic inches of material bringing the material cost up to $12,018. Production time with 0.020” nozzle will be 460 hours.

The tool has 1725 cubic inches of material bringing the material cost up to $12,018. Production time with 0.020” nozzle will be 460 hours. An industrial FDM system capable of manufacturing this mold will cost approximately $350K with $35K/year service contracts. Assume that the system will be in service for 5 years with a $100K trade in value and a 4% interest rate, the machine will have a monthly cost of $7800. With a machine utilization of 50%, the hourly cost of running the machine, excluding overhead associated with utilities and space charges is $23/hr. Therefore, conservatively assume the total hourly charge rate for an industrial FDM system is $40/hr. This would bring the operating cost of the machine to $18,400 and a total part production cost to $31,368.

This would bring the operating cost of the machine to $18,400 and a total part production cost to $31,368. This is still a very attractive cost for composite tooling but highlights how manufacturing time (460 hours for FDM vs 9 hours for BAAM) and low feedstock cost ($8/lb versus $180/lb) can result in a 17X reduction of AM costs ($31,368 for FDM vs. $1870 for BAAM) and 50X increase in production rate.

$1870 for BAAM) and 50X increase in production rate. Figure 6 through Figure 8 show the cost distribution between SLS, FDM and BAAM.Table 4: FDM Tool Cost BreakdownPreprocessing Material Processing Post processing (labor) - Labor - machine TotalTime/material 2 hours 65 lbs 460 hours2 hours 5 hours 9 hoursCost $100 $12,018 $18,400$100 $750 $31,368Percentage 3.2 38.3 58.7 2.71181SLSFDMBAAMPre ProcessingPre ProcessingPre ProcessingMaterialMaterialMaterialProcessingProcessingProcessingPost ProcessingPost ProcessingPost ProcessingFigure 6: Cost distribution of SLSFigure 7: Cost distribution on FDMFigure 8: Cost distribution on BAAMConclusionsWhile additive manufacturing has tremendous potential, current performance metrics in size, speed and cost limit applications. One of the primary cost elements in AM is time on the machine.

One of the primary cost elements in AM is time on the machine. Recent efforts in polymer extrusion show that changing from fiber to reinforced pellets can significantly increase production rate and part size while simultaneously reducing cost. As an example, we show how BAAM can result in a 17X reduction in manufacturing cost and 50X reduction in manufacturing time for the composite tooling industry. Future work is exploring other tooling applications such as stretchform, hydroforming and stamping. Each process will have material performance requirements.AcknowledgementsResearch sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle, LLC. In addition, we wish to show our appreciation to our partners at Cincinnati Incorporated, The Boeing Company, Ford Motor Company, Techmer Applied Materials, BASF and TruDesign.References Holshouser, C., C. Newell, S. Palas, L.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEconomies of Complexity of 3D Printed Sand Molds for Casting Ashley Martof 1 , Ram Gullapalli 2 , Jon Kelly 3 , Allison Rea 3 , Brandon Lamoncha 4 , Jason M. Walker 3 , Brett Conner 3 , Eric MacDonald 31 Youngstown Business Incubator, Youngstown, OH 44555 2 Case Western Reserve University, Cleveland, OH 44106 3 Advanced Manufacturing Research Center, Youngstown State University, OH 44555 4 Humtown Products, Columbiana, OH 44408Abstract: Additive Manufacturing (more commonly referred to as 3D printing) is resulting in a metamorphosis of the sand casting industry as 3D printed sand molds enable castings of unmatched geometric complexity.

The manifold benefits of these molds include: (1) the integration of structural elements such as periodic lattices in order to optimize weight versus strength; (2) the structural inclusion of unique features such as embossed part numbers and/or other details of the production history; and (3) complex geometries that generate new casting applications not possible previously. Additive Manufacturing is often described as providing “complexity for free”, which may not be entirely precise but generally holds true and the identification of castings that are sufficiently complex to benefit from 3D printing is generally left to the intuition of the designer or foundry. New software tools are necessary for foundries to discover opportunities in which the additional costs of 3D printing are compensated by the benefits of increased structural complexity.

New software tools are necessary for foundries to discover opportunities in which the additional costs of 3D printing are compensated by the benefits of increased structural complexity. This paper describes a complexity evaluation tool that scores CAD models to determine the most economical casting approach based on slicing and 2D geometry evaluation. The three potential outcomes include (1) traditional sand casting, (2) AM­enabled sand casting and (3) a hybrid of the two with 3D printed cores in traditional casting flasks. Several case studies are described and evaluated.1.0 IntroductionSand casting is an economical metal process that was used as early as the Shang Dynasty of China in 1,000 BC (Barnard 1961) . In comparison to other forms of casting, sand casting provides a wide range of casting sizes, complexity, and processable metals.

In comparison to other forms of casting, sand casting provides a wide range of casting sizes, complexity, and processable metals. Consequently, the size of the global market recently exceeded 100 million metric tons (Staff 2016) with sand as the predominant method of casting (USITC 2005) . The significance of this ancient industry requires the leveraging of the latest simulation and analysis tools to identify opportunities for 3D printing of both sand molds and cores.Additive manufacturing was invented in the 1980s and was generally relegated to creating form and fit prototypes; however, the possibility of leveraging 3D printing for casting was recognized as soon as a decade later (Hull et al. 1995; Cheah et al. 2004) . The ability to fabricate117geometrically complex parts using layerwise deposition as opposed to traditional subtractive manufacturing now allows for molds and cores that would be otherwise impossible without 3D printing (Bassoli et al. 2007).

2007) . Important benefits of 3D printed sand molds are complex cavities, tight dimensional accuracies, components insertion within the casting and/or mold (ceramics or other metal structures), and an increased freedom in the design of the molten metal delivery system (gating and risers). At least initially, 3D printed sand molds and cores also provide improved delivery times as hard tooling is not required and all molds are rendered from computer designed files in software. Consequently, 3D printed sand molds have lead to an immediate impact on low volume and bridge production as well as design and prototype iteration.Understanding the decision boundary between when a casting is more suitable for traditional methods versus 3D printing is important in order to close the business case.

Furthermore, some scenarios are best suited for a hybrid approaches which allow for complex internal cavities to be formed with 3D printing while the outer features are created using traditional cope and drag casting methodology. Fig. 1 left shows a casting that could not be fabricated with traditional casting while the right figure shows a traditional casting using a pattern and core.Figure 1: Complex casting with 3D printed mold (left) and traditional casting with traditional cores and molds (right ­ courtesy of wikicommons)As a result of the potential benefits, many groups have explored evaluating geometrical complexity in the context of 3D printing for casting in which different challenges were addressed or the benefits of 3D printing were identified (Atzeni and Salmi 2012; Upadhyay, Sivarupan, and El Mansori 2017; Hackney and Wooldridge 2017; Deng et al. 2018; Shangguan et al. 2017; Snelling et al. 2013; J. Walker et al. 2018; Mun et al. 2015; Singh 2010).

2018; Shangguan et al. 2017; Snelling et al. 2013; J. Walker et al. 2018; Mun et al. 2015; Singh 2010) . Research has been conducted on the economics of conventional manufacturing versus additive manufacturing of end usable part production ( Atzeni and Salmi 2012 ; Manogharan, Wysk, and Harrysson 2016 ;118Hopkinson and Dicknes 2003 ; Baumers et al. 2016/1 ; Almaghariz et al. 2016; Conner et al. 2014; Martof and Conner 2017) ; (Thomas and Gilbert 2014) , research has concluded parts with high complexity and/or low volume result in cost savings when utilizing additive for part production. Research on part shape complexity and the effect complexity has on energy consumption in additive manufacturing has also been looked into (Baumers 2012; Psarra and Grajewski 2001) . However, fewer groups have investigated evaluating the complexity of the structure directly in order to inform the economic decisions of using 3D printing for casting (Almaghariz et al. 2016; Conner et al.

2016; Conner et al. 2014; Martof and Conner 2017) . Research has been conducted on the complexity of castings calculated using a regression analysis based on parameters from a CAD model (Almaghariz et al. 2016; Joshi and Ravi 2010) . Part volume ratio, area ratio, number of cores and core volume ratio are a few of the parameters used to calculate the shape complexity. This complexity factor formula was further researched in (Gullapalli 2016) and used to calculate cost of different manufacturing processes. Four different scenarios related to sand casting were considered. Traditional manufacturing of sand casting, 3D sand printing of molds and cores, 3d sand printed cores with traditional molds, and 3d printed patterns. Two cases studies were used to study the relationship between complexity to manufacturing costs and determine the economical manufacturing method.

Two cases studies were used to study the relationship between complexity to manufacturing costs and determine the economical manufacturing method. This previous work, case studies and complexity factor serve as a basis for the algorithms developed in this paper.2.0 Materials and Methods2.1 Complexity AlgorithmComplexity can be calculated with many geometrical methodologies in order to identify the complex features that are relevant to castings, and more specifically, features that require 3D printing (e.g. overhangs, internal voids requiring cores, low draft, ribs, parting line avoidance, cross section uniformity, etc.). These pathological features requiring the use of 3D printing are being indirectly measured in this work by slicing an STL of a final part into layers and then performing analysis on each 2D slice by interrogating detected contours. In this study, four algorithms were identified and were inspired from the works of others (Joshi and Ravi 2010 ; Almaghariz et al.

In this study, four algorithms were identified and were inspired from the works of others (Joshi and Ravi 2010 ; Almaghariz et al. 2016 ; Gullapalli 2016; Baumers 2012 ; Watson 2011) ; (Psarra and Grajewski 2001) and based on the intuition of the authors. For each layer, two complexity numbers were calculated for each algorithm in order to simultaneously evaluate the exterior and interior complexity. Interior complexity is posited to provide data about the need for 3D printed cores, while the exterior complexity informs the decision about whether a flask is required to be 3D printed or if traditional techniques are more suitable.Four algorithms were developed that all began by slicing each of the benchmark STL files and performing analysis per layer producing an average complexity across all layers.

This process was repeated for three orientations of each part and the results were averaged including: an unrotated case, a rotated case 90 degrees in X axis and rotated 90 degrees in Y axis. The119rotations were completed in order to detect a complexity bias relative to orientation. For each orientation, the complexity numbers from all layers were individually calculated and then the total was divided by the number of layers for a mean value that was independent of the number of slices selected. By using more layers, the accuracy was expected to improve by improving the statistical sampling; however each additional layer meant an increase in the duration of computation.

By using more layers, the accuracy was expected to improve by improving the statistical sampling; however each additional layer meant an increase in the duration of computation. For this study, 10 layers were used and each layer image was confined to 1000 rows and 1000 columns ­ all three numbers of which directly impact the computation duration.Each algorithm leveraged existing analysis features of the Python libraries ­ OpenCV and Numpy­STL ­ for finding contours, calculating perimeter length and area and identifying concavity and convexity features of the contour.

OpenCV is open source and when coupled with the libraries Numpy and Matplotlib with Python as a language, provides a framework similar to Matlab and is well suited for efficiently reading STL files and performing a rich set of geometry functions to extract features and evaluate complexity.Algorithm A: Number of contours per layerThe first algorithm was the simplest and just summed the number of contours detected for each layer. Contours that were parents in the hierarchy (exterior contours) were added to the exterior complexity number, and conversely, the child hierarchy elements were counted separately as the interior. Both values were reported to help to identify the correct casting methodology. Fig. 2 shows algorithm A in action with the STL rendered and two layers illustrated.Figure 2: Example of two layers in which contours are summed per layer.

Fig. 2 shows algorithm A in action with the STL rendered and two layers illustrated.Figure 2: Example of two layers in which contours are summed per layer. The STL of a gyroid matrix (left), one layer with nine contours and a score of nine (center) and another example layer with three contours and score of three (right).Algorithm B: Ratio of contour perimeter to contour area per layerAlgorithm B is similar to A except that instead of incrementing the sum of contours, a ratio was calculated and added to a running sum. This algorithm is similar to research conducted by NASA120on perimetric complexity of binary images (Watson 2011). The ratio captured included the length of the perimeter of a given contour divided by the area of the contour. A circle would provide intuitively the smallest ratio and is therefore considered “simple”.

A circle would provide intuitively the smallest ratio and is therefore considered “simple”. An alternative and more complex shape like a star would have a larger perimeter while maintaining a similar area and the increased ratio would emphasize that the shape was more complex than the circle. All ratios per layer were summed and all layers were averaged.Algorithm C: Number of concave features per contour per layerThis algorithm summed the number of contours but also included a sum of concavity defects that were sufficiently concave. A convex hull is the contour that encloses a contour in question but without any concave features ­ the same contour if no concavity exists or a larger contour circumscribing the original. Any point at which the two contours diverge with the greatest separation is defined as a concavity defect and several are shown as red dots in fig. 3 for two different layers of a dragon STL.

Any point at which the two contours diverge with the greatest separation is defined as a concavity defect and several are shown as red dots in fig. 3 for two different layers of a dragon STL. This algorithm provides a score that is directly proportional to the number of contours plus the number of concavity defects for all contours on a given layer. Intuitively, concavity defects are an indication that the structure is complex in the context of the casting as these features can interfere with the use of patterns in traditional casting.Figure 3: Example of two layers in which contours and concavity defects are summed.

The STL (left), one layer with five concavity defects, one contour and a score of six (center) and another example layer with three contours and four concavity defects and score of seven (right)Aggregate Algorithm (D): “All of the Above”This algorithm was an aggregate of the other algorithms: the contours, the ratio of perimeter over area and the number of concavity defects summed together. The aggregate algorithm is the most comprehensive measurement ­ capturing the number of contours, the complexity and concavity of each contour. This consensus approach provided manifold perspectives in order to identify cases that may be pathological in terms of determining if a casting requires 3D printing and was identified as the most effective.1212.2 Benchmark Casting FamiliesSixteen structures intended for casting were selected.

The first two were very simple geometries: a solid sphere and a sphere containing a spherical cavity in order to evaluate the lower bound of complexity from both the internal and external perspectives. In order to provide examples in the upper bound of complexity, a Voronoi tessellation chess piece and a gyroid matrix were selected (fig. 4 left). The queen chess piece was designed using voronoi tessellation. The Schoen’s gyroid (fig. 4. left center) falls into a class of mathematical surfaces referred to as triply periodic minimal surfaces (TPMS) which separate 3D space into two complex and intertwined but separate phases . Here, one phase is established as a solid network of struts, while the second phase is void and constitutes the porous volume. Due to their mathematical complexion, TPMS­derived structures are extremely manipulable and computationally efficient; however they are difficult if not impossible to cast with traditional methodologies.

From a metal casting standpoint, both the chess piece and gyroid part would require 3D printed tooling to cast.Two additional industrially­relevant castings were included as both models were used in a previous economic and complexity study, the results of which are used as a baseline in this study with computer vision libraries. Both castings are provided by Humtown Products with one case including an air brake (fig 4. center right). The design package consists of one mold design and eight core designs. For this research, a family of castings was created with increasing levels of complexity by incrementally increasing the number of cores while maintaining the outer mold line dimensions. Hence a total of eight sub­cases were created by beginning with the first sub­case (a solid casting with one mold and no cores) and then incrementally adding one extra core.

Hence a total of eight sub­cases were created by beginning with the first sub­case (a solid casting with one mold and no cores) and then incrementally adding one extra core. It should be noted that for the complete design, conventional manufacturing involves fabricating eight hard core boxes to make eight cores which leads to issues with damage during transportation, breakage during assembly, misassembly, core shift, and flash at core bonding points. With 3D sand printing, the eight cores become consolidated to a single core.The second previously studied casting was a turbocharger also provided by Humtown Products. This consists of a mold and core design (fig. 4 right) and the complexity was changed by varying the number of cores.

This consists of a mold and core design (fig. 4 right) and the complexity was changed by varying the number of cores. Hence four sub­cases were created, where each sub­case has an additional core and the four cases are: (1) solid turbocharger, no interior cavity and no cores, (2) a single cube shaped cavity (one core), (3) two cylindrical shaped cavities (two cores), and (4) the original design (requiring three traditionally manufactured cores which can be consolidated to one core using 3D sand printing).122Figure 4: Montage of the benchmark structures used in this analysis: (left) voronoi chess piece, (left center) gyroid matrix, (right center) air brake, (right) turbocharger.3.0 Results and DiscussionThe different castings were run through the complexity software and the results are in Table 1 which summarizes the scores of the four complexity algorithms on the sixteen test structures.

These data are then compared to complexity scores assigned by previous work in which the scores were generated manually and correlated well with the decision as to which casting process was most suitable. One major difference with the software automated approach described in this paper is that the original work did not capture internal versus external complexity. To compare the current work to the previous study, a composite score was generated and this was in good agreement with the original work (monotonically related).3.1 Software Complexity EvaluationSelected to demonstrate the lower bound of complexity, the Solid Sphere and Sphere with Cavity both have relatively low exterior complexity valuations, as would be expected. Furthermore, the Solid Sphere has 0.000 or null interior complexity as there are no interior features. Meanwhile, the Sphere with Cavity has low but existent interior complexity corresponding to its simple spherical cavity.

Meanwhile, the Sphere with Cavity has low but existent interior complexity corresponding to its simple spherical cavity. In all cases (except null), the complexity value increases from Algorithm A to Algorithm D, as more features are included in the computation and the difference between the two sphere exteriors is primarily driven by resolution and rounding error. Fig. 5 displays the complexity factor values for each of the benchmark castings as computed by each of the four algorithms for both interior and exterior features.The Air Brake and Turbocharger families further illuminate the ability of the algorithms to discriminate complexity as it relates to selecting a casting technology.. In the Air Brake family, all of the designs have the same external geometry and have increasingly more complicated internal geometries from Design One to Design Eight. The identical exterior complexity of all eight Air Brakes is captured well with all of the exterior complexity scores.

The identical exterior complexity of all eight Air Brakes is captured well with all of the exterior complexity scores. The only minor123differences are calculated and these can be attributed to slightly different geometric boundaries where different internal features interface with the external surface and modified the outer form slightly. Air Brake ­ 1 (AB1), the simplest of all, was determined to have a exterior complexity (for the rest of this paper, all complexity scores are based on the D Algorithm) of 3.128 while AB8, the most complex of all, was determined to have a similar exterior complexity of 2.953. However, while AB1 only had an internal complexity of 1.556, each subsequent Air Brake design had increasing internal complexity culminating in an internal complexity for AB8 of 6.187 ­ nearly four times higher. A similar trend is seen with the Turbocharger family.

A similar trend is seen with the Turbocharger family. These complexity scores, therefore, provide a very efficient method to differentiate between complexity of structures, even when they may appear to be very geometrically similar from the exterior.The Voronoi Chess Piece and Gyroid Matrix are immensely different but provide very clear complexity scores. By visual analysis alone, it is obvious that both benchmark structures have high exterior complexity and the scores capture this. With exterior complexity scores of 17.082 (Voronoi) and 27.580 (Gyroid), the two structures are rated as orders of magnitude more to their exterior surfaces than all of the other benchmark castings. complex with respect Interestingly, both of these structures have extremely low or null interior complexity scores. This is a function of the character of lattice structures. Because all of the interior space is interconnected, little or none of it is discriminated as actually internal.

This is a function of the character of lattice structures. Because all of the interior space is interconnected, little or none of it is discriminated as actually internal. In cases like these, however, the internal score is irrelevant though as the external score is high enough to indicate the necessity for the entire mold package to be 3D printed.Furthermore, differentiating between external and internal complexity is paramount to selecting the most efficient manufacturing method for mold production. Low external complexity generally indicates that a traditional sand mold can be effectively used with hard tooling, whereas high external complexity would likely indicate the need for a complete 3D printed mold package (i.e. The Voronoi Chess Piece and Gyroid Matrix). In the case of low external complexity, though, there are still cases of low and high internal complexity.

The Voronoi Chess Piece and Gyroid Matrix). In the case of low external complexity, though, there are still cases of low and high internal complexity. Low complexity scores for both internal and external may be indicative of a case for entirely traditional methods of mold production ­ that is, cases requiring simple cores or no cores at all. Conversely, the case of low external complexity and high internal complexity would indicate that traditional molding may be used for the cope and drag but additive manufacturing would be effective to address the need for intricate cores.124Table 1: Complexity results for the benchmark structures.125Figure 5: The complexity factor values for Algorithms A­D are shown for each of the benchmark castings. E indicates exterior. I indicates interior.3.2 Comparison to Previous StudiesPrevious complexity factor analysis of the air brake family was completed in (Almaghariz et al.

E indicates exterior. I indicates interior.3.2 Comparison to Previous StudiesPrevious complexity factor analysis of the air brake family was completed in (Almaghariz et al. 2016) using a complexity factor model designed for castings developed by Joshi and Ravi (Joshi and Ravi 2010) . The J&R model requires knowledge of casting design by requiring (among many factors) inputs for the number of cores required and the volume of those cores. A goal of this layerwise complexity factor shown in this paper is to eliminate that design knowledge constraint. Figure 6 shows each of the eight layerwise complexity factors plotted against the J&R model complexity factor for the air brake family of castings. It should be noted that the J&R model considers the casting as a whole and thus includes the exterior and interior complexity together without separation.

It should be noted that the J&R model considers the casting as a whole and thus includes the exterior and interior complexity together without separation. Consequently, exterior complexity factors (AE through DE) are nearly constant compared to J&R model values.A more constructive approach would be examining the summation of exterior and interior complexities. Here we choose the summation of Algorithm D’s exterior complexity (DE) and126interior complexity (DI). This is now plotted against the J&R model complexity factor in fig. 7. The turbocharger family of castings are included in this graph. The trends are now very similar between the two models but the results are different for the two families. A common reference would be desirable. Weighting the layerwise complexity factors is an option.By inspection, one can choose weights and obtain the following relationship (eq. 1):Weighted Layerwise Complexity F actor (W LCF )=0.75 DE+12.5 DIThis is now plotted in fig.

1):Weighted Layerwise Complexity F actor (W LCF )=0.75 DE+12.5 DIThis is now plotted in fig. 8. Here the air brake and turbocharger values lie nearly on top of each other and agree well. In fact, the following polynomial function (eq. 2) can be plotted through the points with a R 2 of 0.9679.WLCF= × 10 6−6JRCF4 − 9 × 10−4JRCF3 + 0.0485× JRCF 2 − 1.2292× JRCF+ 16.243Consequently, the Weighted Layerwise Complexity Factor is a viable method for geometric complexity analysis that does not require detailed knowledge of conventional/traditional fabrication processes.Figure 6: Each of the eight layerwise complexity factor values are plotted against the values calculated in (Almaghariz et al. 2016) for the J&R model.127eq. 1eq.

2016) for the J&R model.127eq. 1eq. 2Figure 7: The exterior (DE) and interior complexity (DI) of Algorithm D are added together and plotted against the J&R model complexity factor.Figure 8: A weighted layerwise complexity factor is plotted against the J&R model128For an economic analysis, we will consider four different methods of fabrication:Traditional manufacturing (TM): Traditional subtractive processes to make molds / cores. ● 3D Sand Printing (3DSP): Complete sand printing of both molds and cores. ● 3D Sand Printed Core (3DSPC): 3D sand printing of cores and conventional pattern making for the mold ­ cope and drag.FDM Pattern­making (FDMP): Conventional core making and 3D printing using fused deposition modeling of hard patterns for conventional mold making with the advantage of faster mold fabrication.The cost model elements for each can be found in (Gullapalli 2016).

Here we consider application of complexity factors and cost modeling to the economics of mold and core sets for the air brake family of castings. Figures 9, 10, and 11 are three graphs plotting cost per part as a function of complexity (using the weighted layerwise complexity factor from above) for three different production quantities: 1, 100, 1000. The process with the lowest cost per part is the most economical for a given value complexity. The most economical process will vary as a function of complexity and quantity. For example, in Figure 9, the most economical process is 3DSP regardless of level of complexity. However, at quantities of 100 as shown in Figure 10, then for castings with complexity of 5 or higher, 3DSPC is the most cost effective method of mold and core sets. For castings of low complexity, then FDMP is the cost effective.

For castings of low complexity, then FDMP is the cost effective. For quantities of 1,000 as shown in Figure 11 then 3DSPC becomes the most cost effective for the full range of complexities for this family of castings. The value of the complexity factor is to help choose the most cost effective method of production.129Figure 9 . Mold and core set cost versus complexity of air brake family of parts for quantity of one.130Figure 10 . Mold and core set cost versus complexity of air brake family of parts for quantity of one hundred.131Figure 11 . Mold and core set cost versus complexity of air brake family of parts for quantity of 1,000.Future work may include the addition of a three dimensional complexity analysis of STL part files. In a previous work, a methodology was developed to calculate 3D signed distance fields of STL files and to subsequently reconstruct 2D and 3D contours based on a level set approach (J. M. Walker et al. 2017).

M. Walker et al. 2017) . The signed distance field generator is agnostic to the input shape and contains a large amount of geometric and topological data including defining internal and external space, distances to all faces, edges, and vertices, and normal vectors of all faces, edges, and vertices. Combined with the layerwise analysis of contours, a supplementary three dimensional analysis score could further elucidate complexity.4.0 Conclusions3D printing stands to transform the casting industry ­ particularly in cases of low volume, bridge In order to inform the business decision as to production, customized or complex castings. whether 3D printing makes sense ­ either fully or hybrid (only cores are printed), this research demonstrated a relatively simple 2D slicing approach for measuring complexity of the casting geometries.

The finding show that of the four algorithms, the “all of the above” gave values that provide a decision boundary that was in line with the intuition of the authors as well as obvious extreme cases (sphere as a case of simplicity or a gyroid matrix for a case that can only be cast132with the help of 3D printing). This layerwise complexity factor was compared to a known complexity factor for conventional casting fabrication method and showed similar results but without requiring design knowledge of the traditional methods. The economics of complexity and quantity were shown for a traditional casting tooling method and then compared to three methods that involved additive manufacturing.AcknowledgementsWe would like to thank the Friedman Endowment for Manufacturing at Youngstown State University for supporting this project.ReferencesAlmaghariz, Eyad S., Brett P. Conner, Lukas Lenner, Ram Gullapalli, Guha P. Manogharan, BrandonLamoncha, and Maureen Fang. 2016.

In this study we examine the potential of using 100% recycled ABS to form filaments for use in Fused Deposition Modelling (FDM) 3D printing. We then characterise the resulting changes in the printing quality and mechanical properties, over a single recycling cycle. We found that ABS can undergo recycling and reforming into consistent printer filaments without the addition of virgin material. However, notable changes in polymer characteristics were observed, reflected by degradation in mechanical properties during tensile tests and a decrease in the polymer melt flow, which required reduced raster speed to achieve repeatable prints.

Despite these limitations, we demonstrate that recycling and reprinting is possible with acceptable loss of material integrity, and could provide unique opportunities for sustainable use of waste ABS using 3D printing technology.1 IntroductionPlastics are a highly pervasive and integral part of our daily life, found in everything from packaging, furniture, automobiles, micro devices and even biomedical components [1-5]. The prevalence of plastics can be attributed to their highly desirable material properties, such as their lightweight, durability and amenability to be manufacturing using a diverse range of standardised processes, such as casting, CNC machining [6], laser ablation/engraving [2], injection moulding [7] and 3D printing [5, 8].

The usefulness of plastic and the optimisation of its use in manufacturing has resulted in an escalation in its global production, rising by over 500% in the last three decades, and is set to increase to 850 million tons annually by the year 2050 [9]. This intensive production and use of plastics is resulting to several unprecedented issues relating to the disposal and recycling of this material given the sheer volume that is to be processed. Additionally, most plastics are not biodegradable and produced from oil.

Additionally, most plastics are not biodegradable and produced from oil. Therefore plastic recycling is becoming increasingly important, not only due to finite fossil based resources on our earth which are simultaneously used for energy generation, but from the resulting negative impact our current actions are have on the environment.532Arguably, sustainable management of plastics has now become one of the biggest challenge we now face in the modern day.In an ideal scenario, the lowest impact on the environment would be the direct re-use of a plastic products and devices. However, this may not always be practical should a product become compromised or damaged beyond the effective usefulness. Additionally, in the case of typical plastic FDM printing, a significant volume of material is lost due to the use of rafts, supports and failed prints, leaving such volumes of material typically redundant.

In such scenarios recycling becomes the most attractive option to make best use of this finite resource.With respect to additive manufacturing, plastics arguable comprise the most widely used feedstock for 3D printers and have become ubiquitous in our daily lives through the popularity and accessibility of polymer FDM printers. FDM systems operate by the use of thermoformable plastic filaments which are progressively fed into a heated nozzle, and the liquefied material is extruded outwards to form the respective layers of a printed object [10]. In essence the filaments used in FDM printers are identical to the filaments produced during the extrusion phase of the classical industrial recycling process. Therefore we believed that using these same processes with waste recyclable plastics, otherwise destined for landfill, we can generate filaments for use in FDM based 3D printers.

Therefore we believed that using these same processes with waste recyclable plastics, otherwise destined for landfill, we can generate filaments for use in FDM based 3D printers. Such concepts are rapidly becoming of research interest to many groups worldwide [9, 11-13].In this study we investigate the reformation of 100% recycled 3D printed ABS material, into 3D printer filaments and assess the printing quality and integrity of this as a primary FDM feedstock. We investigate the parameters relating to the production of FDM filament using both virgin pelletised and recycled granules and characterise the resulting filament diameter and melt flow rate for varying extrusion temperatures. Using the most promising filament extrusion parameters and resulting filament, we examine the print quality in terms of layer adhesion and surface roughness of the printed material. We also perform tensile testing of printed samples to assess material degradation resulting from the recycling process.

We also perform tensile testing of printed samples to assess material degradation resulting from the recycling process. Our initial findings confirm that repeatable FDM 3D printer filaments can be produced using 100% recycled ABS and used to print parts with a similar build quality to commercial printer filaments, yielding near equal surface roughness profiles. We found there was a marked decrease of 13 to 49% in the ultimate strength of several test samples, depending on the print orientation. Overall, we have found that the superficial print quality and ultimate strength to be acceptable for most printing applications, leading us to believe that recycling of ABS plastics could provide a viable means of upcycling waste 3D printed ABS material.

This technique could help both research and home users obtain greater longevity from 3D printer feed stocks and potentially reduce the escalating burden on landfill.2 Methodology2.1 Plastic Granulation ABS plastic was obtain from either the use of virgin pellet material (Filastruder, GA, USA) or from failed/redundant 3D prints and raft/support material from teaching classes and research use within the School of Engineering at Deakin University. No restrictions were imposed to ensure that the recycled ABS can from a single filament supplier source or ABS grade, with the exception of the ABS pellets, which were all sourced from the single batch. This generic use of ABS for recycling purposes ensured that there was no bias in the results, whereby plastics containing specific plasticisers or additives which could enhance in the recycling process were utilised.

Therefore our results are likely533to be reflective of general conditions experienced when recycling batches of ABS for wider reproducibility of our findings.Waste ABS was granulated using a combination of crushing in a bench vice, direct cutting using wire cutters and the use of modified paper shredder. The bench vice was used crush larger printed models into smaller pieces (approximately >8cm3) by clamping and applying torque to compress the respective model. The wire cutters were used to perform finer size reductions, before the resulting pieced were fed into the modified paper shredder to produce granules of ABS. To ensure a relatively uniform size of granule, pieces were separated using a sieve with a mesh size of approximately 5mm.2.2 Filament generation To produce the 3D printer filaments, both the pelletised material and the granulated ABS was processed through a commercial filament extruder (Filabot Ex2, Barre VT, USA).

The extruder operates by the loading of plastic material into a hopper, which then using a drive screw, delivers material through a temperature controlled heated region. At sufficient temperature, the plastic softens and liquefies before being forced out of the device through a nozzle of a fixed diameter due to growing pressure from the continuous addition of plastic material. The diameter of the resulting filament can be controlled based on the temperature of the heating zone, the rotational speed of the drive screw and the size of the granules placed into the hopper and the nozzle diameter. The temperature effectively controls the viscosity of the loaded ABS once it enters its molten state, while the drive screw rotational speed and granule size influence the feed rate, which in turn dictates the pressure applied to the molten material for extrusion.

Combined, these factors allow for control of the ABS melt flow rate out of a fixed size extruder nozzle, which then dictates the resulting diameter of the extruded filament. By careful control of all of these parameters, we are able control the filament diameter to obtain filaments of a suitable size for use in the 3D printers. The printer employed in this study was designed to accept filaments with either a diameter of 1.75mm or 2.85mm and could readily accept filament diameters between this ranges. For this initial study, we opted to produce 1.75mm filaments. Owing to the lack of any robust spooling equipment, to generate the filaments we utilised a simple gravity extrusion based methodology. In this process the extruded material was allowed to flow from the nozzle, at a height of 860mm, and gradually make its way to the floor under its own weight.

In this process the extruded material was allowed to flow from the nozzle, at a height of 860mm, and gradually make its way to the floor under its own weight. We found that this methodology worked effectively for the use of ABS to realise a simple solution for continuous production.Filaments were generated using virgin pelletised ABS and 100% recycling ABS. For each variant of ABS, care was taken to ensure a relatively uniform granule size. When producing the filament, the extrusion speed was fixed at 50% (approximately 17.5rpm), the nozzle diameter at 1.75mm, and the temperature adjusted accordingly to obtain filaments of approximately 1.75mm.

Once filament production had reached a steady state, for each temperature and ABS variant we recorded the length of material that was extruded over a one minute time period to estimate the melt flow rate of the polymer.2.3 3D printing and characterisation All 3D printing in this study was performed using a commercially available FDM printer (Flashforge Dreamer, USA), which utilises Flashprint proprietary software interface and allows for control of all printing parameters (extrusion temperature, feed rate, etc) to optimise the final models. The printer also has a heated bed to allow for improved adhesion of ABS during the building process. For this study, the virgin filament material was printed using standard printing parameters, comprising534235oC nozzle temperature and 105oC bed temperature respectively.

For this study, the virgin filament material was printed using standard printing parameters, comprising534235oC nozzle temperature and 105oC bed temperature respectively. For the recycled material, the bed temperature was held constant unless issues with adhesion were observed, and other influential printing parameters were adjusted stepwise until optimal model production was achieved (no gaps, even layer adhesion, etc).Figure 01: An overview of the ABS recycling process, comprising 1) ABS recovery and granulation 2) filament extrusion and spooling, 3) FDM 3D printing and 4) the resulting final models.To examine the material properties of the recycled ABS, as compared to virgin ABS, various test specimens for tensile tests were printed accordance to ISO 527-2 European standards for plastics [14].

Given limited availability of recyclable ABS plastics for this study and in keeping in the spirit of the sustainable credentials of this study, test specimens were made to the type 1BA standard, which minimised overall material consumption. Tensile tests were performed using an Instron 5960 Dual Column 50kN tensile test system (MA, USA), where five repeat measurements were conducted. Only results which conformed to the standardised test procedure were recorded and all other samples were discarded until five measurements were achieved.

Only results which conformed to the standardised test procedure were recorded and all other samples were discarded until five measurements were achieved. Beyond tensile testing, various test models were fabricated to qualitatively examine the printability of the recycled filaments in realistic scenarios, for ‘passive’ (models) and ‘active’ parts (nut and bolt) to demonstrate the viability of using 100% recycled ABS material.3 Results3.1 Granulation With respect to commercial filament generation, the industry standard comprises the processing of highly uniform pelletised material in a melt extrusion process. We adopt a similar535methodology for the generation of recycled plastic filaments, however make uses of coarse granulated material, attained by repeated grinding in the modified paper shredder. Twenty virgin ABS pellets were randomly collected and their diameters measured.

Twenty virgin ABS pellets were randomly collected and their diameters measured. The average pellet size was determined to be approximately 3.4±0.4mm, and ranged across diameters of approximately 3-4.5mm. This value formed the baseline average we would aim to achieve with the granulated recycled ABS material, thereby ensuring that this variables influence on the polymer feed rate was kept uniform between the recycled and commercial ABS.Typically, the waste plastic required 4-5 repeated shredding cycles and processing through the 5mm mesh sieve to achieve suitable granules. Figure 1 shows a sample of the granulated material obtained from the shredding and sieving process. Twenty randomly selected granules were measured for their average diameter. The average granule size was determined to be approximately 3.8±1mm, and ranged across diameters/maximum lengths from between 2.3-6.3mm, in the examined samples.

The average granule size was determined to be approximately 3.8±1mm, and ranged across diameters/maximum lengths from between 2.3-6.3mm, in the examined samples. It was found that despite being generally effective, the current methodology did at times allow for larger granules to pass through the sieve. Standard methods for palletisation of the recycled material would have undoubtedly increased the granule diameter consistency. In this study we opted against this approach as to do so would result in potential thermal degradation of the recycled polymer material, which in turn would influence the subsequent mechanical tests. Ultimately, the obtained average would allow for a relatively uniform material feed rate into the extruder.3.2 Filaments To produce the filaments, either the pelletised or granulated plastics were loaded into the hopper of the extruder and fed into the heating barrel for production at a fixed extrusion speed of approximately 17.5rpm.

From preliminary tests we established that extrusion was possible at approximately 180oC, however showed signs of containing coarse polymer material. This temperature was used as a baseline lower limit for extrusion and working in steps of approximately 5oC, the average diameter of the filament was measured up to a maximum of 220oC. For each temperature setting the melt flow characteristics of the polymer were realised.It was found that there was a measureable difference in the filament diameter until the extrusion process has reached a steady state (with the material touching the floor and coiling). We found that there was an approximate difference of approximately up to 0.4mm in the measured diameter, and believe to be due to the gravitational forces of the filament stretching under the weight of the previously extruded material.

Once the steady state had been reached, the gravitational force remained relatively constant due to an approximate equal mass of material being between the nozzle and floor, resulting in a relatively uniform filament diameter for further extrusion. At this point a measured maximum deviation of 0.15mm was found from the average filament diameter.Figure 2bi) illustrates graphs of the change in extruder filament diameter and the average extrusion speed, for changes in the temperature for both the pelletised and 100% recycled ABS material. With respect to filament diameter, it can be seen that both ABS variants show a decrease in the average filament diameter for an increase in the extrusion temperature. This would be expected as at the start temperature of 185oC, which is well above the glass transition temperature of ABS (≈105oC), the polymer is in a liquid phase state.

This would be expected as at the start temperature of 185oC, which is well above the glass transition temperature of ABS (≈105oC), the polymer is in a liquid phase state. Therefore increases in temperature will result in a decreasing the relative viscosity of the bulk polymer, resulting in greater melt flow through the nozzle (as observed in figure 2bii), and hence a decrease in the overall extrusion diameter. Interestingly, for 100% recycled ABS, a relatively uniform melt flow rate from the extruder was observed, with a modest536rise of approximately 5-6mm3/s from temperatures of 185 to 205oC, as can be seen in figure 2bii). This discrepancy could be explained by the irregularities in the average granule size of the recycled ABS, which results in unpredictable fluctuations in the feed rate of the ABS into the extrusion nozzle, and hence sporadic changes in the melt flow rate.

This thought is substantiated by the relatively uniform increase in melt flow rate for the pelletised ABS, which has a much greater uniformity in average pellet size, and hence more predictable feed rate characteristics. However, there is exists a possibility that the relatively uniform melt flow rate could be due to some form of degradation of the ABS material. To test this assumption, should we perform prints with increasing nozzle temperature, we should expect the extrusion volume to remain relatively constant. We therefore conducted test prints varying the nozzle temperature to check this hypothesis, as detailed in section 3.3.Figure 02: a) Illustration of the recycling melt extrusion process, whereby the recycled ABS granules are loaded into the extruder and the resulting filament is gravity fed to the floor.

b) Graphs illustrating the results for i) change in filament diameter for increased extrusion temperature and ii) Change in melt flow rate for increased extrusion temperature.Overall, examining the results in figure 2bi), the pelletised virgin and 100% recycled ABS show similar rates of diameter decrease for a respective rise in temperature, excluding the spurious result at 210oC. It was found that to obtain a filament diameters of 1.75mm, the virgin pelletised ABS was to be extruded at approximately 205oC and the recycled ABS at 200oC respectively. Speculating, this decrease in the extrusion temperature may be indicative of a decrease in the mechanical integrity of the ABS resulting from thermo-oxidative degradation, as has been described previously [15, 16].

Using the resulting extrusion temperatures filaments of 1.76±0.12mm we produced using the pelletised ABS and 1.76±0.14mm for the recycled ABS respectively for subsequent mechanical testing and qualitative printing analysis.537Figure 03: a) Graphs illustrating the surface roughness of printed section of a square test print using both commercial and recycled 3D printer filaments. b) Pictures of the square test print and surface topography profiles when using i) commercial and ii) recycling 3D printer filaments. c) Test print structures from the recycled ABS filaments comprising i) a bolt and 3 nuts and ii) a propeller concept.5383.3 3D printingUpon formation of the filament from each respective ABS variant, the resulting material was loaded into the 3D printer for testing.

Examining the filament produced using the pelletised ABS, it was found that the printing performance matched that of standard commercial ABS filaments, with optimal printing parameters comprising nozzle temperature 235oC, print speed of 60mm/s and bed temperature of 105oC. When the recycled ABS was printed at these same parameters, it was found that the resulting prints showed signs of under extrusion in the form of gaps in the extruded strands of the print (results not show). Typically, to remedy this issue using commercial filaments, one can raise the nozzle temperature, which results in an increase in the melt flow rate of the ABS and a greater volume of deposited material. As we recorded a relatively uniform melt flow rate for the recycled material, an increase in the nozzle temperature would not result in an increase in the deposited material. This was confirmed by some preliminary printing tests examining nozzle temperatures of 240, 245 and 250oC.

This was confirmed by some preliminary printing tests examining nozzle temperatures of 240, 245 and 250oC. In this instance to compensate for the under extrusion, the printing speed was reduced, with optimal printing parameters comprising nozzle temperature 235oC, print speed of 50mm/s and bed temperature of 105oC. It is also speculated that we could equally increase the filament feed rate to compensate for the under extrusion, and we hope to examine this in future studies. Ultimately, we achieved reliable printing parameters to achieve consistent printed models and so began to scrutinise parameters relating to the printing quality and material integrity.3.3.1 Print quality testsTo examine the printability of the recycled material in comparison to virgin material we printed a test structure for comparative purposes, comprising a square of 20x20x5mm, and performed various surface profile measurements.

Figure 3 illustrates the final test structures and the resulting 3D topography scans, alongside data for the surface roughness over a region randomly selected region following the direction of the deposited strand. It was found that the average deviation in the surface was relatively uniform between the virgin and recycled ABS material, with an approximate deviation of 50µm between the peaks of trough of where the deposited ABS material join. Equally the average surface roughness of the deposited material was found to be similar, with an average roughness along the length of a deposited ABS stand of 2.08µm for commercial filament and 2.16µm for recycled filaments respectively. These test would indicate a similar print quality between the commercial and recycled filaments for the derived printing parameters.

These test would indicate a similar print quality between the commercial and recycled filaments for the derived printing parameters. To qualitatively demonstrate the print quality of the recycled filament, we performed test prints of designed engineering items, comprising a simple nut and bolt system and a propeller concept. Both items were challenging prints for even commercial grade filaments and were printed successfully in recycled material on our first attempts, as illustrated in figure 3c).3.3.2 Tensile testsAs we had now established that the printing quality of the recycled filaments could match the performance of commercial filament, we examined the changes in the mechanical properties of the printed material as a metric to examine material degradation. Tensile tests were performed by539clamping a test specimen in the Instron test rig (figure 4ai), while a constant load was applied beyond initial elongation and on to failure.

Tensile tests were performed by539clamping a test specimen in the Instron test rig (figure 4ai), while a constant load was applied beyond initial elongation and on to failure. Samples were printed all three potential print orientations, as seen in figure 4aii), using a print infill of 100%. Sample results from the tensile tests can be seen in the graphs in figure 4b), which illustrates a comparison between the commercial and recycled printed specimens for samples manufactured in the X build orientation. A summary of the complete test data can be seen in table 1.It was found that there was minimal difference in the ultimate strength and average stiffness for use of the virgin ABS filament in the X and Y build orientations, and an approximate 25% comparative decrease for samples build in the Z orientation.

By comparison there was a clear decrease in the recorded ultimate strength and average stiffness for the recycled ABS, across all build orientations, with the X build orientation offering the greatest level of mechanical strength and rigidity. Comparing the ultimate strength between the virgin and recycled filaments, it can be seen there is a clear decrease of between 13 to 49%, with the X build orientation demonstrating the lowest percentage decrease (13%). Equally, there was a marked decrease in the material stiffness with measured decreases of between 17 to 28%, with again the X build orientation demonstrating the lowest percentage decrease (17%). Clearly these results are an indication of the mechanical degradation of the ABS plastic during recycling, but also hint that the build integrity of the test sample (e.g. layer adhesion) is also compromised for the given parameters utilised.

layer adhesion) is also compromised for the given parameters utilised. Should this be the case, it would explain why the results for the ultimate strength would be greatest for the X build orientation and lowest for the Z build orientations respectively given the translational mode of the printer nozzle during the polymer extrusion process.Figure 04: a) i) Photograph of the tensile test apparatus, ii) Diagram of the examined build orientations for the printed dog bone samples and iii) printed dog bone samples awaiting testing. b) Graphs depicting the results of tensile tests for dog bones built in the X build orientation and 100% infill for i) commercial and ii) recycling 3D printer filaments.540In terms of the strain failure of the test samples, there was marginal difference between the commercial and recycled samples, with the exception of the samples in the z orientation, which showed an approximate 50% decrease.

We believe these results are again indicative of the compromised layer adhesion of the recycled material, which would be exacerbated for samples manufactured in the Z orientation. Indeed, it was visually observed that there was compromised layer adhesion on the Z build samples prior to tensile tests.Results ultimately confirm that there is degradation in the mechanical properties of the recycled material, but also that there may be issues with the manufacturing process and the sample type utilised. In a recent study by Torrado et al 2016, varying ultimate tensile strength (UTS) was observed between samples built in differing ASTM standard tensile types, rastering direction and orientation [17]. There findings would perhaps indicate the there are issues with relying on FDM printed samples built vertically (Z build orientation) to determine the UTS, particularly between the varying size types as part of recognised standards.

We believe their study highlights the necessity to validate our claims in this study with further tests using tensile samples of increased size and control of the rastering direction. We hope to perform further tests in future continuations of this study.SampleX Build Y Build Z BuildUltimate Strength (MPa) 40.59±1.42 42.34±0.88 33.25±0.66Commercial Filament Average Stiffness (N/mm)430.50 419.78 404.05Strain at Failure (%)10.35±1.73 6.62±0.28 4.62±0.51Ultimate Strength (MPa) 35.44±2.01 29.06±1.21 16.81±0.66Recycled Filament Average Stiffness (N/mm) 357.40 302 329.80Strain at Failure (%)11.62±1.13 6.36±0.18 2.47±0.23Table 01: A summary of the tensile tests for each of the examined printing orientations using both commercial and recycled filaments.4 ConclusionsIn this study we have demonstrated a methodology for the reforming of 100% waste ABS plastics into relatively uniform 3D printer filaments (1.76±0.14mm).

When forming the filaments we found that increasing the extrusion temperature reduced the extruded filament diameter through the extruder die. Interestingly, the melt flow rate of the recycled material remained relatively uniform for increased extrusion temperature, whereas virgin pelletised ABS showed an increase in melt flow rate. This was suspected to be due to thermal degradation of the recycled ABS, and required the adjustment of the raster speed to avoid under extrusion when printing. It was found that the superficial print quality of the recycled filament was of a similar consistency as commercial filaments, as confirmed by surface roughness analysis. However there was a marked degradation in the mechanical properties of the recycled ABS, resulting in a 13 to 49% decrease in the ultimate strength of the printed material.

However there was a marked degradation in the mechanical properties of the recycled ABS, resulting in a 13 to 49% decrease in the ultimate strength of the printed material. Ultimately, we believe that our methodology to extrude 100% recycled ABS plastic to be a viable option to make better use of printed resources, and could prove useful in the management of other sources of waste ABS, in addition to alternative recyclable plastics.5415 ReferencesR. C. Thompson, S. H. Swan, C. J. Moore, and F. S. vom Saal, "Our plastic age," Philosophical Transactions of the Royal Society B: Biological Sciences, vol. 364, pp. 1973-1976, 2009. M. I. Mohammed, M. N. H. Z. Alam, A. Kouzani, and I. Gibson, "Fabrication of microfluidic devices: improvement of surface quality of CO2 laser machined poly (methylmethacrylate) polymer," Journal of Micromechanics and Microengineering, vol. 27, p. 015021, 2016. B. Luijsterburg and H.

A. Boldizar and K. Möller, "Degradation of ABS during repeated processing and accelerated ageing," Polymer Degradation and Stability, vol. 81, pp. 359-366, 2003/01/01/ 2003. A. R. Torrado and D. A. Roberson, "Failure Analysis and Anisotropy Evaluation of 3D-Printed Tensile Test Specimens of Different Geometries and Print Raster Patterns," Journal of Failure Analysis and Prevention, vol. 16, pp. 154-164, 2016// 2016.[1][2][3][4][5][6][7] [8][9][10][11][12][14][15][16][17]542.

M TOOLING BY SOLID FREEFORM FABRICATION & ELECTROPLATING TECHNIQUESR. Ippolito, L. Iuliano, Politecnico di Torino (ITALY) A. Gatto, Universita di Ancona (ITALY)Abstract The term "Rapid Tooling" refers to the rapid creation of tools in much the same way as Solid Freeform Fabrication (SFF) means the rapid creation of models. This revolutionary approach offers both designers and manufacturers attractive advantages in the form of time compression and cost reduction. Timesaving is of vital significance in the production of EDM electrodes for the fabrication of moulds and dies. The employment of SFF and electroplating for this purpose is described. The performances of these new type of electrodes were compared with those obtained conventionally in terms of dimensional tolerances, surface roughness and working time.

The performances of these new type of electrodes were compared with those obtained conventionally in terms of dimensional tolerances, surface roughness and working time. The surfaces ofboth electrodes and workpieces were also examined by SEM microscope.1.IntroductionConstant improvements in the performance of Solid Freeform Fabrication (SFF), also well known as Rapid Prototyping (RP) technologies [1,2,3] and both terms will be used in the paper, have greatly extended their market. This is particularly true with regard to Rapid Tooling (RT), the name given to the fabrication of tools used to make small runs of parts for construction of the prototypes of mass-produced objects.

This is particularly true with regard to Rapid Tooling (RT), the name given to the fabrication of tools used to make small runs of parts for construction of the prototypes of mass-produced objects. Both the technical and the commercial literature, in fact, show that other undeniably interesting applications are overshadowed by rapid tooling.A sector of major importance that has so far received little attention is the fabrication of electrodischarge machining (EDM) electrodes, especially those employed in the manufacture of small and medium-sized dies and moulds. The application of RP techniques in the direct mould manufacture by metallisation [4] has some limits: •it is fairly easy to use only for the production of simple plastic injection moulds and probably for sheet metal drawing dies; it can be employed to make simple dies without the moving parts needed to make the undercuts.

The fabrication of an EDM electrode by SFF, on the other hand, would be followed by unrestricted conventional manufacture of the die or mould. The economic aspects of this combination have still to be determined.2.Production of EDM electrodes by Solid Freeform FabricationThe RP model [5] itself could, in principle, be used as the electrode, or employed as a master for the creation of the final electrodes through further processes. The first alternative is unfeasible, since SFF techniques cannot be easily used for the direct production of metal objects, and even if the present limitations were overcome the metal structure would be unsuitable for an electrode. The second requires further operations, but results in good precision and an excellent finish.Three methods can be used to obtain an electrode from a RP negative master:copper spraying: this has proved feasible in a number of tests.

Its drawback is that electrode wear is considerable during machining due to the porosity of the copper;199investment casting: this method has proved satisfactory in other applications. Electrodes made in this way, .however, are unsuitable for EDM, since their structure results in marked granular corrosion and rapid deterioration;electroplating: the RP model is coated with a conductive lacquer and clad with a layer of copper in a galvanic bath. This method is inexpensive, since the forming operation is unmanned. In addition, the copper is very pure and dense. The long processing time required and some geometrical limits, however, are the drawbacks.A-A30C/.J26,4~C/.J34,420Figure 1: Negative master of the EDM electrode.Electroplating offers the best trade-off between advantages and disadvantages.

In the light of the current capacities of RP techniques [6,7] in terms of dimensional tolerance and surface finish, it may be supposed that the electrodes produced in this way will only be suitable for roughing operations.2-Experimental workThe feasibility of constructing a electroplated EDM electrode was initially investigated with a part whose geometry was relatively complex (fig. 1), yet readily measurable with a co-ordinate measuring machine (CMM). Electroplating was carried out in an experimental cell (fig. 2) designed and built for the purpose at the University of Ancona. It has an acid bath and a soluble copper anode, and is equipped with: • a temperature control system to keep the bath at 23°C, (range of recommended temp.:[20+30rC);a fumes exhaust system; • a controlled voltage D.C.

supply unit connected to a continuity set, since if copper were deposited chemically during a power failure it could prevent the build-up of the plating layer when the current comes on again. The process parameter values have been chosen to maximise the growth rate of the layer onceit is a few millimetres thick.200Continuity unitFumes ejectionPump & heater controlHeaterCompressed air- . - . . IHeat exchangerFigure 2: Sketch of the Cu electrodeposition cell.The type of conductive paint applied on the surface of the negative master is one of the factors that determine the quality of the result, since it has a great influence on the rate of growth and the dimensional precision. Preliminary tests were run with copper and silver charged paints and with gold sputtering. A silver paint was chosen on the basis of four parameters: ease of application, covering power, growth rate resolution.

A silver paint was chosen on the basis of four parameters: ease of application, covering power, growth rate resolution. The master was built by stereolithography (SLA), as this widely used technique ensures good resolution and a good surface finish [7].and geometricalThe first question to be settled was the dimensional stability of the master in the galvanic bath, since one of the main problems with polymers is their hygroscopic behaviour. The master was formed on a 3D System machine with the new CIBATOOL SL5170 epoxy resins. A wax master was also made with a Desktop NC manufacturing system.Both masters were covered with an approx. 3 mm layer of copper. Insertion of a metal stem and filling of the shell with resin to stiffen the structure obtained by electroplating and lock the stem inside it completed construction of the electrode.

Insertion of a metal stem and filling of the shell with resin to stiffen the structure obtained by electroplating and lock the stem inside it completed construction of the electrode. The dimensional tolerance, the shape errors (flatness, roundness and cylindricity) and both horizontal (~) and vertical (-t) surface roughness of the two masters and electrodes were measured and the results are illustrated in Table I.The precision of the two masters is much the same and the electrodes obtained from them are of practically the same quality. In view of this demonstration of the stability of the SLA and the fact that it showed no signs of alteration and could therefore be used again after removal of the copper shell, it was decided to make two more electrodes: • one from the SLA master as before; •the other by conventional milling with an NC machine tool. Both electrodes were then used to make a cavity in a block of AISI 1040 carbon steel.

Both electrodes were then used to make a cavity in a block of AISI 1040 carbon steel. TheSLA master, copper electrode and workpiece are shown in fig.3.201Table I: Master and electrode dimension, shape and roughness measurements.Element Max nShape error (Ilm)Flatness 9 34 3 16Roundness C15 14 17 144 4 5.4 5.4[*] n: number of tolerance unit (5).Figure 3: a) SLA master, b) eu electrode, c) workpiece; both electrode and workpiece were sectioned to make SEM observations. On section of electrode is visible the deposed layer of copper.The EDM parameter values were set for each test for evaluation of electrode performance. Dimension, shape and surface roughness measurements were taken from the masters, electrodes and workpieces. Both electrodes and the machined workpieces were also examined at SEM .As can be seen in Table III, the dimension and shape tolerances of the two electrodes are of the same order of magnitude (see also fig. 4).

4). There are two reasons for this: the conventional milled electrode was made to wide machining tolerances for roughing operations; the intrinsic difficulty of obtaining tighter tolerances when working on a malleable material such as electrolytic copper.4.Results and discussion4.1 Machining performanceEDM is very slow when compared with conventional material removal processes. As much material is possible is usually removed by conventional drilling and milling to save time and minimise electrode wear. This was not done in our study since the aim was to assess tool performance under the hardest conditions202A commercial EDM machine was used, with hydrocarbon as the dielectric. The test parameters were: voltage (V), the discharge time (t), and the efficiency ratio (t), defined as the ratio between t and the duration of the test sequence (discharge plus idle time).

The parameter values are set out in Table II. The electrode obtained from the SLA master used more current at the same voltage. Its removal rate (Vas), however, was faster than that of the conventional electrode.Table II: EDM test parameter values.ConventionalSFF (SLA)First stage t [ms] V [V] I [A] t [tip] Vas [mm3/min] t [h] Second stage t [ms] V [V] I [A] t [tip] Vas [mm3/min] t[h]300 =45 7 0.9 13 2.2750 =45 18 0.9 255 3.9300 =45 14 0.9 18 1.6750 =45 20 0.9 302 3.34.2Geometrical and roughness measurementsThe geometrical and shape tolerances of the masters, electrodes and workpieces were determined with a CMM, which measures parts at all points with the same nominal dimensions. The surface roughness of the parts was evaluated with a Hommel T 1000 Stylus instrument using a sampling length of 4.8mm.

The surface roughness of the parts was evaluated with a Hommel T 1000 Stylus instrument using a sampling length of 4.8mm. As regard the dimensional error our evaluation model [7] introduces the number of tolerance unit (n) for 90% of the observations as the quality index, It is included in the third column of Table ill and is also illustrated in fig.4. The shape errors (flatness, roughness and cylindricity) are listed in the forth, fifth, and sixth columns, and the roughness measurements in the seventh and eighth columns.As can be seen, there are no significant differences in the maximum tolerance grade and the roughness measurements of the workpieces.

The worse flatness and cylindricity of the part machined with the electrode derived from the SLA master can probably be attributed to the difficulty in its assembly.It can therefore be asserted that the technology of galvanic deposition of copper in thickness on models constructed by SLA can be used with success for the production of EDM electrodes to be employed in roughing operations.4.3 After machining the surfaces of both electrode, derived by SLA master and the workpiece were sectioned, cleaned ultrasonically and observed at SEM microscope.SEM observations203Table In: Comparison of the results of the EDM electrode manufacturing methods.MethodElementElectrode before machiningMaxn 90% 422Flatness 5Ra(ffi.

Roundness Cylindricity ~ 3.Shape error (mm).724Conventional Electrode after45687164.14.4SFF (SLA)machining Workpiece RPmodel Electrode before machining Electrode after machining Workpiece480 240 20422051820 3 1668411 17 14162024 41 26256816 5.4 5.46.6. 1413 6.2 6.27.113The shallow craters, roundish remolten parts and cracks (fig.5) reported for EDM of varioustypes of steel [8] were observed on the workpiece.In the horizontal zone of the hole, the globular parts are present in greater quantity with respect to the vertical surface (fig.6); this fact is imputable to the different efficiency of washing on the inside of the hole during the machining. A similar difference was noted on the electrode in the form of a greater iron contamination (see the arrows in fig.7) from the workpiece on the horizontal as opposed to the vertical part.

A similar difference was noted on the electrode in the form of a greater iron contamination (see the arrows in fig.7) from the workpiece on the horizontal as opposed to the vertical part. Reciprocal tool-workpiece contamination has also been reported for EDM operations with conventional tools [8]. The outer white layer and the altered layer immediately below it are similar to those observed by other authors [9].WorkpieceElectrode before machining600518n2000...&....--ConventionalSFF (SLA)Figure 4: Comparison of both electrodes and workpieces in terms on number of tolerance unit forthe 90% of observations.204Figure 5: SEM image of the vertical surface of the holeFigure 6:SEM image of the vertical (a) and horizontal (b) surface of the workpiece.5.ConclusionsFour conclusions can be drawn from this study:1. the method proposed gives electrolytic copper electrodes from SLA masters.

the method proposed gives electrolytic copper electrodes from SLA masters. Other SFF techniques can not be used, because they do not allow an acceptable surface finish to be obtained;2. the dimensional precision of these electrodes is on a par with that of those produced conventionally for the same type of operations;3. the deep thickness obtainable suggests that a 0.5 - 1 mm machining allowance can be provided for conventional finishing, which would mean that the method could equally be used to produce EDM electrodes for surface finishing operations;4.

the performance of these electrodes is decidedly superior to that of a tool obtained by mechanical machining from a bar nominally of the same materiaL205Figure 7:SEM image of the horizontal (a) and vertical (b) surfaces of the electrode.No assessment can yet be made of the economic aspects of this method, since its feasibilityrequires further study.Endurance tests are now in progress to determine the behaviour of the new electrodes over time, together with trials of the method in the fabrication of an injection mould. Their completion will allow initial conclusions to be drawn concerning the time and cost required for the production of an electrode of average complexity.6.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEDUCATION OF ADDITIVE MANUFACTURING- AN ATTEMPT TO INSPIRE RESEARCHLi Yang*Department of Industrial Engineering, University of Louisville, Louisville, KY 40292AbstractAlthough additive manufacturing technologies have undergone significant development in recent years, significant challenges remain in the understanding of the physics of the processes as well as many other aspects. Therefore, in the education of the next generation AM workforce, beside the instruction of existing AM knowledge, it is also of critical importance that the state-of- the-art research subjects and concepts are made aware to the students.

In various classroom and lab setups at University of Louisville, contemporary AM research subjects are introduced to students via various tools including self-guided literature review, mini-research projects and design competitions. Through literature review based studies the students not only become aware of various research subjects, but also have the opportunity to practice critical analysis skills with new AM knowledge.

Through literature review based studies the students not only become aware of various research subjects, but also have the opportunity to practice critical analysis skills with new AM knowledge. Through the mini-research and the design competition project-based learning processes, the students not only gain hands-on experiences with AM practice but also learn about the research methodologies employed with various AM research subjects.Introduction First introduced in late 1980s, the Additive Manufacturing (AM) technologies have evolved into a large group of manufacturing technologies that are radically changing the horizon of the manufacturing industries via transformative capabilities in accommodating high levels of design complexity, design customization and production flexibility [1]. Over the past decade, AM has grown by over 50 times in terms of market size, and has maintained a growth rate of over 18% in the past 4 years [2].

Over the past decade, AM has grown by over 50 times in terms of market size, and has maintained a growth rate of over 18% in the past 4 years [2]. Meanwhile, the applications of AM have also significantly expanded from predominantly prototyping, tooling and fixtures to many other functional components in high value-added industries such as biomedicine, aerospace, automobile, electronics and energy [2]. In many of these functional applications, the most sought-after capabilities of AM is the enabling of Design for Functionality (DFF), which focuses on maximizing the functions and performance of the structures with minimum resource consumptions (e.g. materials, production time, defect rate, etc.). This in turn requires that the users of the AM technologies are fully informed of the design capabilities and constraints introduced by AM. However, it is also broadly perceived that such knowledge base is currently not well established [3, 4].

However, it is also broadly perceived that such knowledge base is currently not well established [3, 4]. Currently the existing knowledge with AM is still heavily empirical, especially in the areas of process development, manufacturability and design optimization [5]. In many application areas, specific knowledge and expertise that are based on individual products or specific material and process combinations usually lacks generality. Although the research and development of AM have gradually shifted towards more systematic investigation of the process physics in recent years, a coherent theory structure for many AM technologies remain an open challenge. Furthermore, the rapid evolution of the AM technologies aggravate this issue by constantly pushing the boundaries of the existing knowledge.

Furthermore, the rapid evolution of the AM technologies aggravate this issue by constantly pushing the boundaries of the existing knowledge. For example, the recent development of the printing technologies for the fiber-reinforced composite materials has significantly disrupted the traditional view that AM lacks capability with this type of materials44[6, 7]. Similarly, the observation that the inert gas flow is a dominant factor in the occurrence of the spatter phenomenon in laser based powder bed fusion processes, which contradicts with the traditional perception that this is driven mostly by the keyhole recoil pressure [8].With these challenges, effective AM education and training using the traditional lecture based learning is difficult. On one hand, a considerable percentage of the knowledge space exists in the ever-evolving horizon, which necessitates the constant update of lecture contents and the presentation of the state-of-the-art research in related areas.

On the other hand, many of the newly generated knowledge are presented in fragmented scientific publications, which lacks systematic structures or even consistency and require extensive literature research. Such obstacle can be overcome by multiple measures. For example, the lecturer can perform literature research periodically and keep the contents updated and structured, or that the lecturers can prompt the students to perform such literature research as part of the learning experiences. The use of literature review as a learning tool has been reported previously in AM education [9], and in the author’s teaching practices similar approach has been taken for the introduction of AM concepts and state- of-the-art applications, which will be introduced in more detail in this chapter.As AM is still under intense development, there still exist many open questions and challenges that likely call upon continuous investigations.

For example, in the investigations about the AM process quality benchmarking, despite various proposals about the benchmark artifact designs, there is still a general lack of consensus with the intended purposes and the effectiveness of these designs [10, 11]. Consequently, when such knowledge aspects are covered in the education processes, it becomes necessary to present it as an open subject. This not only makes the students aware of these contemporary issues but also encourages them to consider the problems from different perspectives. The exposure to these subjects in turn ensures that the next-generation workforce for AM can continue to dedicate to these critical issues and eventually provide solutions to them. There exist various pedagogical methods that facilities the critical thinking and active learning of specific subjects, such as debate based learning [12], project-based learning [13], group discussion [14], and problem-based learning [15].

Among these methods, the project-based learning can be effectively utilized to also implement hands-on learning with AM processes, which is a critical part of the AM education since many of the required skills for AM professions are closely related to the manufacturing operations. The project-based learning is commonly employed in AM education in many educational institutes in order to implement hands-on exercises, however in most cases the learning objectives focus on the comprehension of the freeform fabrication design concept enabled by AM layer-wise processes. Further development of teaching strategies are needed to effectively combine the existing strategies with the unique challenges of the AM education.

Further development of teaching strategies are needed to effectively combine the existing strategies with the unique challenges of the AM education. In this chapter, some attempts in utilizing the project-based learning for the introduction of open questions of AM technologies are described in an effort to inspire future works in developing AM teaching methodologies.Literature review based learning The literature review based learning was employed in an introductory AM course in the Department of Industrial Engineering at University of Louisville. The basic concept of this method was previously described in detail [9]. The course consists of three components: the lectures, the student presentation and a semester project paper. Literature review based learning was employed45as a major tool for the self-guided learning of specific subjects related to the AM technologies and their applications. The two parts of the semester project are elaborated in more detail below:1.

The two parts of the semester project are elaborated in more detail below:1. Presentation: the students will give a 30-45min presentation about the selected subject of literature review. The format and style of the presentation are generally unrestricted. However, the students were required to prepare the presentation as a “lecture”, which will be consequently considered as the materials for the test.2. Paper: the students will submit the literature review papers towards the end of the semester (about week 11-12), which will be reviewed and commented. The reviewed papers will be returned to the students for revisions, which mimics the journal article review process. The final version of the paper will be submitted at the end of the semester (about week 15 or 16). At the beginning of the semester, the students were provided with the detailed instructions of the literature review project in order to encourage early preparations.

At the beginning of the semester, the students were provided with the detailed instructions of the literature review project in order to encourage early preparations. A sample list of subject was provided as reference, which is partially shown in Table 1. The subjects ranging from AM applications (e.g. AM applications in dentistry) to specific state-of-the-art AM technology development (e.g. 4D printing) and some contemporary issues related to AM (e.g. AM cybersecurity). The students were encouraged to combine their personal or other research interests with the literature review and choose their own subjects. On the other hand, the students were required to submit a preliminary outline for the literature review sub-topics within two weeks of their initial subject identification in order to receive feedbacks regarding the scope of the works. This is because that the scopes of some of the subjects (e.g.

This is because that the scopes of some of the subjects (e.g. AM in aerospace applications) could potentially be too broad and must be refined to fit into the scope of semester projects. The students work in groups of 2-3 members and proceed with the finalized subject afterwards.During the first half of the course, lectures were given for specific AM technologies following the ASTM F2792 standards, which include material extrusion, vat photopolymerization, powder bed fusion, material jetting, binder jetting, directed energy deposition, and sheet lamination [16]. In addition, direct write technologies were also introduced as a separate category in order to give the students some perspectives about the multi-disciplinary AM systems. These lectures were expected to provide the students with fundamental understanding of the AM technologies that would allow them to perform more informed literature review works.

These lectures were expected to provide the students with fundamental understanding of the AM technologies that would allow them to perform more informed literature review works. Various aspects about general AM technologies, such as process physics, material compatibilities, advantages and disadvantages in practice, and current application areas, are introduced in the lectures. The contents of these lectures must be reviewed and updated by the lecturer every time the course is offered in order to ensure most up-to-date information.During the second half of the semester, each student group would give the lecture presentation about the subject they chose for literature review. Due to the setup of such assignment as a “lecture”, the students must ensure that the technical details are presented in clear and coherent ways.

Due to the setup of such assignment as a “lecture”, the students must ensure that the technical details are presented in clear and coherent ways. In addition, the presenters would be asked to also emphasize on the interactions with the audiences as well as the Q&A sessions, which was found to often significantly improve the engagement levels of student audiences for better learning outcomes. The lecturer would use the matrix listed in Table 2 for the evaluation of the presentations and determine whether the presented themes and information would be adopted for test questions. It is important that the students are46aware of such process so that they would dedicate more considerations about the communication aspects with their presentations.

Furthermore, such peer learning approach aims to achieve multiple objectives beside the communication of specific knowledge, including collaborative work, critical enquiry, communication and articulation of knowledge, self-management and self- assessment, which are also critical skill for many contemporary engineering working environments [17].Subjects about AM applicationsSubjects about new AM technologiesSubjects about contemporary AM issuesAM in medical applications such as bone implants, surgical planning, dentistry, external prosthetics, and tissue engineering; AM in soft and hard tooling; AM in foundry applications; AM in aerospace applications; AM in energy applications; AM in architecture and construction; AM in food industries; 4D printing; Organ printing/bioprinting; Multi-material or functionally graded material printing; Finishing technologies for AM; AM supply chain issues; AM cybersecurity;Table 1 Sample subjects for AM literature review projectOrganization and content (20%)  Individual presentation organization/flow (good intro, transition flow)  Use of visual aids (demonstrations, graphics, etc.)  Use of time (fit naturally within 14-16 min) Presence (15%)  Physical appearance (Adequate for lecture)  Posture, gesture, movement, eye contact  Enthusiasm, enunciation, clarify (no monotone, not too quiet) Delivery and grammar (25%) Knowledge of materials and terminology (evidence of depth of knowledge) Overall effectiveness of delivery method (ability to connect with audience) Freedom from distracting “Uh”s, etc.

(no nervous words or motions) Overall  Clear thesis statement and purpose (what is the main point?)  Adequate support of thesis (how did you back up the main point)  Overall material flow and organization (easy to follow, logical, smooth(Graded by individual) 1 2 3 4 5 1 2 3 4 5 2 4 6 8 10 (Graded by individual) 1 2 3 4 5 1 2 3 4 5 1 2 3 4 5 (Graded by individual) 2 4 6 8 10 2 4 6 8 10 1 2 3 4 5 (Graded by group) 1 2 3 4 5 1 2 3 4 5transitions)2 4 6 8 10Definite conclusions (summarize main points)  Overall completeness of topic coverage (any gaps in coverage?)  Q&A session-knowledge of topic1 2 3 4 5 2 4 6 8 10 1 2 3 4 5Table 2 Grading criteria for the literature review presentationFor the literature review paper, the students were required to follow the author’s guidance of popular AM journals such as Progress in Additive Manufacturing (Springer International Publishing AG, ISSN 2363-9512), Rapid Prototyping Journal (Emerald Publishing Limited, ISSN 1355-2546), Additive Manufacturing (Elsevier, ISSN 2214-8604) and 3D Printing and Additive Manufacturing (Mary Ann Libert Inc, ISSN 2329-7662).

Each group would be required to include at least 20 journal publications, 10 conference proceeding publications, and 5 industrial articles47(e.g. trade articles, company press articles) as the citation for their papers. As it was recognized that many students might not have adequate background for literature search, a lecture session was dedicated to the introduction of scientific publication database searching. Various literature search tools including Google Scholar, ScienceDirect, conference proceeding archives and article references were introduced to the students. The students would also have the options of actually submitting their papers to the journals if the subjects are deemed appropriate.

The students would also have the options of actually submitting their papers to the journals if the subjects are deemed appropriate. On the other hand, regardless of whether these papers are eventually submitted as real journal manuscripts, the students would have the opportunities to enhance not only the understanding of the specific subjects but also various other skills, including technical writing, structured thinking, engineering problem identification and critical thinking. A sample grading rubric is shown in Table 3, which aims to provide the students with some guidelines about the paper writing. In Table 3, the technical review contents only consists of 25% of the grades. This is because that many of the technical content issues would be addressed in the initial review feedback.

In Table 3, the technical review contents only consists of 25% of the grades. This is because that many of the technical content issues would be addressed in the initial review feedback. For literature review, the organization of materials and the structure of the paper are also of critical importance, and therefore are also emphasized (15% grade).The use of literature review based learning was welcomed by the students, and many reported that they not only gained in-depth understanding with the AM technologies but also became more proficient with technical writing. Due to the broad range of subject selection, the students could usually identify ones that are of most interest to them, and therefore exhibit high level of motivation and engagement throughout the project.

Some of the frontier subjects such as 4D printing and organ printing are constantly of interests to many students, which would give not only the review group but also the entire class the opportunity of in-depth learning of the subjects.Draft/final paper completed and submitted on-time  Good abstract and keywords  Followed formatting rules/appearance  Appropriate use of references  Appropriate use of figures  Grammar and spelling  Effective use of case studies/examples  Effective thesis/support/conclusions  Flow of paper/organization  Completeness of content15% 5% 5% 5% 5% 5% 10% 10% 15% 25%Table 3 Grading rubric for literature reviewMini-research projects Various open questions in the AM research and development fields can be potentially developed into project based learning materials.

In the development of these projects, additional factors that should also be considered include the hardware capability of the educational facility, scope of project and the estimated time consumption. For example, for subjects related to the metal powder bed fusion AM process principles, the lecturer must make decisions such as whether to set up hands-on experiments using the often expensive equipment or computationally costly finite element simulation based studies. In addition, many of these open research questions would likely require additional refinement by the lecturer so that a subset of the problems can be set up as48student projects. For example, in the discussions of the standard benchmark designs of AM, it was suggested that the benchmark designs could serve as geometry benchmark, property benchmark and process development benchmark [11].

As the simultaneous design of multiple benchmark types has been proved challenging, when setting this up as a learning project, one of the benchmark types such as geometry benchmark could be set up as the objective.In the Design for Additive Manufacturing (DFAM) course also at the Department of Industrial Engineering at University of Louisville, various hands-on lab projects were set up as listed in Table 4. These projects attempt to cover a range of DFAM subjects from process capability investigation and material process-property characterization to lightweight geometry designs. The projects were also designed in a sequence that would facilitate progressive learning about the integrated DFAM theory. In the first lab project, the students were asked to use AM to design and realize a relatively simple real-world product with certain geometrical and mechanical requirements.

In the first lab project, the students were asked to use AM to design and realize a relatively simple real-world product with certain geometrical and mechanical requirements. Using the desktop material extrusion processes which has lower process resolutions, the students were set up to perform an “impossible” task, which would demonstrate the limitations of the AM product realization and the need of DFAM concepts. The second lab project aimed to help the students to establish the process geometrical quality design concept for AM, and was combined with the open research subject of AM geometric benchmark design. The students were given a short lecture about the current state-of-the-art research with the geometry benchmark designs and their design strategies.

The students were given a short lecture about the current state-of-the-art research with the geometry benchmark designs and their design strategies. The students were then presented with various objectives with the geometry benchmark designs such as the characterization method for Geometry Dimensioning and Tolerancing (GD&T) with AM parts, considerations of variability with AM processes, and easiness of metrology. The students worked in groups of 2-3 and propose benchmark designs to the classes in competition for the final selection. The selected design would then be fabricated using various AM technologies (e.g. powder bed fusion, vat photopolymerization, material extrusion, binder jetting) by individual groups and consequently used for metrology.

powder bed fusion, vat photopolymerization, material extrusion, binder jetting) by individual groups and consequently used for metrology. Fig.1 shows some of the benchmark designs proposed by students in the previous semesters.Lab sequence123Lab project themeUsing AM for product realizationAM Geometric benchmarkingFDM process-property evaluationDescription Students would realize a product design with geometrical accuracy/strength requirements, such as a medical syringe or phone microscope A geometric benchmark part design designed by students would be fabricated using various AM processes.

The geometrical characteristics of each processes would be evaluated and compared Students perform experimental designs for FDM system in order to evaluate the effect of process parameters (temperature, scan speed, part orientation, part size, etc.) on mechanical properties of the structures45Geometrical design evaluation for cellular structure Design efficiency of lightweight structureCellular structures with various unit cell designs will be fabricated and tested in order to compare their propertiesComparing the topology optimization design and the unit cell design and evaluate their efficiency and manufacturabilityTable 4 Hands-on project for project based learning49Fig. 1 Student designs for AM geometry benchmarkingFollowing the completion of the first two lab projects, the students would have opportunity to gain some additional in-depth understanding about the design for mechanical properties with the AM processes.

This lab would focus on the introduction of the geometry-process-property design for the AM technologies, which is different from many traditional manufacturing processes that often allows for the separate design of geometries and material properties. This project was usually set up by using the material extrusion processes since the fabrication costs are relatively low. The design of material properties for the material extrusion structures include various aspects, such as the infill pattern, the number of skins/shells, the printing parameters (speed, temperature, bed preheating, etc.), the part orientation, the part dimensions, and the use of slicer software. In order to minimize the unnecessary implications induced by the software slicing algorithm, the project could be set up on a single printer. This also reduces printer-to-printer variability if such design factor is not intended to be investigated.

This also reduces printer-to-printer variability if such design factor is not intended to be investigated. Table 5 shows the experimental design table used for the investigation of mechanical properties of “standard” tensile coupons, while Fig.2 illustrates another project setup that focuses on the investigation of the effects of infill designs on the mechanical properties of the printed parts.Printing temperature(°C) Printing travel speed (mm/s) Angle of tilt (°) Coupon cross sectional size (mm) Layer thickness (mm) Number of shells Infill density Raft205, 215, 225 100, 150 0, 90 3, 5, 7 0.3 2 20% YesTable 5 Example experimental design for mechanical property evaluations of material extrusion AMFig. 2 Project example for infill pattern effect for material extrusion AM50The students would be normally given about 2-3 weeks to work on this project depending on the schedule and availability of the printers.

Due to the large size of the experimental designs, the project was split into smaller experimental groups and assigned to each groups. Consequently, after collecting the testing results, the students would be asked to submit the data to the lecturer, who would compile the results and send them back to the class for the data analysis and final reporting.In the last lab project, the students would focus on the design of AM lightweight structures. This subject has been of great interest for the AM research and development communities, therefore the focus of the project was to expose the students with different design methodology of AM lightweight structures, including the topology optimization and cellular designs. Depending on the AM platform intended to be used for the project, the level of complexity of the lightweight designs can be pre-defined to ensure manufacturability.

Depending on the AM platform intended to be used for the project, the level of complexity of the lightweight designs can be pre-defined to ensure manufacturability. For example, when material extrusion printers are used, the topology optimization designs could be restricted to 2D, and the cellular designs could be restricted to 2.5D extruded lattices such as honeycomb (Fig.3a). On the other hand, when laser sintering powder bed fusion processes are employed, the geometry design restrictions could be largely removed.ZY2 x u n i t c el la. 2D designs (material extrusion)b. 3D designs (laser sintering powder bed fusion)Fig. 3 AM lightweight design project examplesThe students were asked to design for lightweight structures for static 3-point bending load for a bar structure using both topology optimization and cellular designs. For the topology optimization designs, commercial topology optimization software (e.g.

For the topology optimization designs, commercial topology optimization software (e.g. solidThinking or Fusion 360) usually provide relatively straightforward topology optimization modules for single-loading cases and could be used by the students for the projects. For the cellular structure designs, the students were instructed to utilize unit cell based design methods, which could be realized via commercial CAD software. The lightweight designs could be evaluated potentially via finite element analysis (FEA), however the students were required to overcome potential practical issues such as file format compatibility issue and computational cost issues.

The setup of this design project closely imitates the current engineering practice, which aims to not only facilitate the learning of the subjects but also give the students the crucial experiences for AM lightweight engineering designs.Design competition The semester project of this course was also setup specifically to stimulate knowledge synthesis and application. The project theme was set up as a lightweight component design for a “real” product, which required the students to combine the knowledge gained from the previous lab projects and to apply optimization synthesize with the knowledge in the design practices. One of such examples of project theme is the car bumper design. As shown in Fig.4a, the students were asked to design the car bumper that would be assembled to a standardized car body. The students51were allowed to use any design methodology (empirical, theoretical, simulation) to achieve the design objectives below:a.

The students51were allowed to use any design methodology (empirical, theoretical, simulation) to achieve the design objectives below:a. Maximum impact energy without damage on the car bumper; b. Maximum impact energy without damage to the passenger.Regular eggs were used as passengers to increase the level of challenges to the designs. As the design objectives are somewhat mutually exclusive and are difficult to achieve simultaneously, the students must utilize the knowledge learned from the course to devise optimized strategies. Such challenges are also currently encountered in the application of AM lightweight designs, which often result in different trade-off scenarios that are arbitrary and empirical. Such diversity in design preferences were also reflected in the course, which is illustrated in Fig.4b by the different designs used by the students.Bumper Assembly Featurea. Semester project design exampleb. Sample designs from studentsFig.

Semester project design exampleb. Sample designs from studentsFig. 4 DFAM course semester design projectTo increase the student engagement to the project, the final evaluation of the design was set up as a competition. As shown in Fig.5, the car-bumper assemblies were placed on a descending slope with egg passengers. The cars would slide down the slope and hit a relatively sharp object placed at the bottom of the slope in order to create the impact event. A short discussion session would be held after the competition to give the students additional opportunities to analyze the car bumper designs and summarize the learning outcomes. After the competition, the students would also be given additional time (1-2 days) to submit a final report with written analysis. By analyzing the final report and comparing it with the report of the first lab project, the learning outcomes of the DFAM theories could be quantified.52Fig.

By analyzing the final report and comparing it with the report of the first lab project, the learning outcomes of the DFAM theories could be quantified.52Fig. 5 Final evaluation competition of the semester projectConclusive mark There exist various challenges and risks in the integration of the state-of-the-art AM knowledge into the teaching practices. One obvious challenge is the relative lack of knowledge structure and sometimes even the validity of some of the conclusions. It becomes more difficult for the students to fully comprehend the problems when the overall understanding is still partial. Therefore, instead of teaching knowledge, a possibly more effective approach is to focus on the instruction of research landscape, which can be more difficult to evaluate outcomes for. In addition, it also results in additional pressures for the lecturers in the preparation.

In addition, it also results in additional pressures for the lecturers in the preparation. There is the risk of mis-instruction, which can result from the incorrect summarization of the literatures or the lack of clear conclusions. To avoid this, the lecturer must be keen to continuously follow the research and development trends in the AM fields.Despite these challenges and risks, it is of crucial importance for the AM education to incorporate state-of-the-art research and development subjects.

With the goal of educating the next-generation workforce in mind, it should be realized that unlike many other well-developed manufacturing engineering principles, AM will likely face rapid development for the coming decades, which makes it critical for the workforce to be aware of the dynamic landscape of the knowledge base and be more engaged in the continuous research and development efforts.Acknowledgement The author is grateful of the partial support of the America Makes/GE Global Research grant #401004534. Many of the educational efforts reported were supported by the Rapid Prototyping Center at University of Louisville.Reference [1] Gibson, I., Rosen, D., Stucker, B. (2015). Additive Manufacturing Technologies: 3D Printing, Rapid Prototyping, and Direct Digital Manufacturing. Springer.53[2] Wohler’s Report 2016. (2016). Fort Collins: Wohler’s Associates. [3] Gao, W., Zhang, Y., Ramananujan, D., Ramani, K., Chen, Y., Williams, C. B., Wang, C. C. L., Shin, Y.

Minimum feature width is approximately 25 microns, and can be reduced further.EFABcan be used to manufacture micromachines and microelectromechanical systems (MEMS), offering significant advantages over current processes: e.g., true 3-D geometry, IC compatibility, low capital investment, and process automation.BackgroundMicrofabrication. During the last decade, there has been extensive worldwide R&D in the field ofmicromachininglmicrofabrication. There is much interest in replacing macroscopic devices with micromachines andmicroelectromechanical systems (MEMS).toreduce size, weight, and cost while often increasing reliability and performance.

There is much interest in replacing macroscopic devices with micromachines andmicroelectromechanical systems (MEMS).toreduce size, weight, and cost while often increasing reliability and performance. There is also a desire to create devices offering entirely new capabilities not possible to implement macroscopically (e.g., surgical instruments, mechanical high resolution displays).Current microfabrication approaches generally fall into three categories [1]: Bulk micromachiningremoves material selectively from regions ofa substrate (sacrificing precious chip "real estate" and producing very simple shapes); surface micromachining (Fig. 1) deposits thin structural layers and sacrificial layers onto a substrate using semiconductor-type processing; and LIGAgenerates 2.5-D "extrusions" (Fig. 1) by molding polymers in metal tooling electroplated into apertures in synchrotron exposed photoresist.

1) by molding polymers in metal tooling electroplated into apertures in synchrotron exposed photoresist. .All these approaches produce simple geometries (e.g., 1-5 differentcross sections), are not standardized or particularly generic (usually requiring a custom process for each new device), and require many manual steps·(each using specialized equipment). Moreover, they usually donotpermit integrated circuitry to be part ofthe device, involve large numbers ofprocess variables (often requiring multiple iterations to perfect), arevery costly (in facilities, capital equipment, labor, and materials), and involvelonglead times.As a manufacturing paradigm, Solid Freeform Fabrication can potentially overcome the shortcomingsofcurrent microfabrication approaches.

SFFcan offer arbitrary 3-D geometry; short leadtirrles;fu1lY automated, unattended processing; a single, self-contained machine thatpr()(Jucesa huge/yariety of products; device cost that is largely independentofcomplexity; highrepeatability (few processing variables); and easy device design (few manufacturing constraints). In.d~d, in.themicro-and mesos¢opic1 University of Southern California, Department of Materials Science & Engineering161(betweenmicroandmacro)domain,wh~thereare no existing generic fabrication technologies(such as millin~ andcastingthat~etakefor~tedin1h,emacroscopiq world), many devices\~ made with great difficulty orRotatall.Here SEE.Cal1 gofar beyond itsusualroleas a tool to increase speed and reduce cost, becorninginsteadatool1hatrnakespossibleforthe.firsttimethe manufactureofa entire class ofnew products.Fig. 1.

1. Typical devices produced by sutface micromachining (left) and LIGA (right).UnfortUllately, most SFF.processes, intend¢d.ias they are for producing;macrosoopic parts, nonnaUy cannotprO<1Fefeatureslessthan 50-100 IJIl1 wide, or layers less than 50-150 IJIl1 thick. Moreover, virtually all SFFprocesses generate alayer serially (one volume element ata time), so are too slow and costly for quantltyProduction. Nor d0ltlaterlaisofi:en allow functionallJ$e,idtletointrinsic properties or porosity. Somepr~ses are not net-shape'requiringStlPpOrt stru~swhichmust be explicitly designed and removedfium each device.

Somepr~ses are not net-shape'requiringStlPpOrt stru~swhichmust be explicitly designed and removedfium each device. For example,mi~ostereolithography[2], while succeeding in generatingsmall features and layers, is limited in throughput, produces parts from photopolymer, and requires supports.ElecfrodepQsitioIlforSFF.lfanSFF process couldbe developed based onselective electrodeposition ofmaterial, it mightbeU$e<l formiCfQ-' andmeso-mbrication, overcomingthe limitations ofexisting mierorabrieati0niandSFF processes. EI~O<1eJ?9siti<>nis a tremendously important manWacturingprocess used to finish m¢talparts(inthe .~. ofeleqtroplatirlg),and.alsoto.manufacture metal parts by deposi~ metal over rnolds (in the case ofelectrof~nning).El~odepositiontakesplacein an electrochemical cell consistingoftwo~lectrO<1es,an anodeand.acathode.(i.e., woIkpiece or mold) and an electrolyte. Compared with the additive approaches used by most SFFprocesses, electrodepositionis fundamentally different.

Compared with the additive approaches used by most SFFprocesses, electrodepositionis fundamentally different. The materlalis "grain-Iess"'de}Jositedone atonliclayer~tatime (vs. particle-by-partic1e or drop-by-drop); and the solid produced can have desirable material properties, since itis fully-dense and relatively homogeneous and isotropic (e.g.,inter..layerjunctions can ~ avoided). Moreover, material may be d~sited over an entire layer (e,g., hundreds or thousands ofparts) simultaneously.An earlierapptoach to SFF by electrodeposition,invented at MIT, arranged for deposition to 0CCtlf preferentially in a localized region dueito highcmtentdensity near a sharpened anOde that is moved in a3-D path [3].

However, even us~multipleanode!),1hisprocess seems too slow for mass.Production, is unable to easilt~tepartswith overhangs (sinceitlacks asupportmaterial), and suffet'S frOll1llnprecisely-defined f~s.Moreconventionalapproachesto selective electrodeposition, apparently never attempted for use in SFF,are''b1UShp~'', inwhiqhelectrolyteisappliedlocally to a substrate; and "laser..enhanced plating", in whichthe plating rate islocally~lerated by heating. These Processes too are essentially serial and therefore slow, and have difficulty producing vet)' small, well-defined features.The standard approach to high-precision selective electr~sition-used in the manufacture ofPC boards and read-write heads--is known as "through~mask plating". Through-mask plating can be used to d~it features 1 IJIl1 or smaller.

Through-mask plating can be used to d~it features 1 IJIl1 or smaller. With this method, the substrate is ~ared, photoresist is applied, the resist162is UV eX(X)sed and developed, the substrate is plated through apertures in the insulating resist, and the resist is finally stripped. This entire process involves a total ofup to nine separate steps, the use ofseveral different machines (often located in a cleanroom), and multiple liquids. Automation would lead to a machine ofgreat complexity and high cost. Furthermore, total cycle time is larg~, so repeating these steps hundreds or thousands oftimes to fabricate a 3-Dpart would be impractical.Instant MaskingA new method ofachieving selective electrodeposition is therefore necessary for a viable SFF process. We have devised such a method, which we call "Instant Masking".

We have devised such a method, which we call "Instant Masking". As with through-mask plating, Instant Masking uses photolithographically-pattemed masks, and deposits material simultaneously over the entire layer (which may consist ofa batch ofdiscrete devices). Yet unlike through-mask plating, Instant Masking allows the photolithographic process to be performed completely separate from the part-building process, and allows all ofthe masks to be generated simultaneouslyprior to part generation, rather than during part generation (again and again for every layer). This separation also makes possible a simple, low-cost, automated, self-contained machine that builds devices, while the photolithographic steps can be handled in a traditional cleanroom (e.g., with masks provided by a service bureau).Anode+--MaskBlanket depsosfted 2nd materialSelectively deposfted 1st material(b)(c)Substrate. . .(d)(e)Fig. 2.

. .(d)(e)Fig. 2. Stages in the EFAB process.Instant Masking selectively depoSits material onto a subsQ."ate (or previous layer) by simply mating a patterned mask against the substrate, depositing material throtlgh apertures in the mask, and then uri-mating the mask. Since the mask may be topologically complex (e.g., consist ofisolated "islands"ofmasking material), it must be attached to a support structure, which in our currentimpletnentation is simplythe anode ofthe electrodeposition cell. This configuration is shown in Fig. 2 (a).163EFABWe call the new SFF process we are developingSacrificial (support) materialbased on selectiveelectrodeposition using InstaJ:l,t Masking-EFAB (Electrochemical FABricatiQn). A completedlayer in EFAB consists ofboth structural and sacrifiqial tilllterial, as shown in Fig. 3.

A completedlayer in EFAB consists ofboth structural and sacrifiqial tilllterial, as shown in Fig. 3. The block of sacrificialmaterial in which EFAB-builtdevices are temporarily embedded serves the same purpose as sacrificialmaterial usedinSFFprocesses such as those ofCubital and Sanders Design: mechanical support ofstructural material. This elim.ihl;ltes most restrictions on device geometry, allowing the structural material on a la)'erto overhang---..andevenbec.lisconnected from---that ofthe previous layer. Such geometrical fteedolllialso makes possible monolithically-fabricated "assemblies" ofdiscrete, interconnected parts (as withCubital's Geneva mechanismdemonstrated in the early '90s). This is a necessity for micro- and mesoscopic devices, for which actual assembly is generally not possible. Furthermore, the placement ofthe sacrificial material, and its ultimate removal in an automated, batch process, is not device-specific (avoiding the need for explicit support design).Fig. 3.

Furthermore, the placement ofthe sacrificial material, and its ultimate removal in an automated, batch process, is not device-specific (avoiding the need for explicit support design).Fig. 3. Plan view ofa layer in the EFABprocess.In our current implementation, the process consists ofthe following steps, repeated on every layer: 1) selectively deposit first material; 2) blanket deposit second material; 3) lap. The process flow is shown in Fig. 2. In Fig. 2 (a), the first material is selectively deposited onto a substrate, producing the (incomplete) layer shown in Fig. 2 (b). In Fig. 2 (c), the second material has been blanket-deposited over this, contacting the substrate in those regions not platedwith the first material. Then, as in Fig. 2 (d), the entire two-material layer hasbeen lapped to achieve precise thickness and planarity. After repetition ofthis process for all layers, the structure shown in Fig. 2 (e) is etched to yield the desired device (Fig.

After repetition ofthis process for all layers, the structure shown in Fig. 2 (e) is etched to yield the desired device (Fig. 2 (f).While some implementations ofEFAB could selectively deposit both structural and sacrificial materialsusing Instant Masking, we are currently employing lapping in the process for several reasons: 1) it allows the use ofblanket deposition ofone material, thus requiring only halfthe masks, and bypassing any alignment ofthese masks to those used for the other material; 2) neither material needs to be deposited with an extremely uniform thickness to achieve high accuracy and good masking; 3) layer thickness-and thus Z accuracy-ean be precisely controlled.All SFF processes require software togenerate cross sectional layer geometry from 3-D CAD geometry. With EFAB, one or more high-resolution (e.g., +/- 1 f.UIl) photomasks must be prepared which include the geometry ofall the cross sections ofthe device.

With EFAB, one or more high-resolution (e.g., +/- 1 f.UIl) photomasks must be prepared which include the geometry ofall the cross sections ofthe device. For now, we generate mask files using layout editor software; however, we are developing custom software which will import .sTL files and export a 2-D photomask file in a format accepted by commercial photomask making equipmentPatterned maskFig. 4. Schematic of first-generation EFAB machine.164ExperimentalMaterials. EFAB can be used to fonn structures from any electrodepositable metal, with the only constraint being that the accompanying sacrificial metal can be selectively removed (e.g., by chemical etching) after the layers are fonned. Nickel and copper fonn a desirable material system, with copper as the sacrificial material, and it is this system that we have focused on thus far. Nickel has good strength, temperature, and corrosion resistance, and its magnetic properties are useful in electromagnetic devices.

Nickel has good strength, temperature, and corrosion resistance, and its magnetic properties are useful in electromagnetic devices. Meanwhile, copper can be selectively etched with respect to nickel. We employed commercially-available electrolytes (i.e., plating baths) such as acid copper, cyanide copper, pyrophosphate copper, Watts bath, and nickel sulfamate.First-generation machine. We constructed the machine shown schematically in Fig. 4 to allow us to investigate the Instant Masking process. This machine uses an interchangeable copper anode, which fonns the floor ofa plating tank. The substrate, a nickel disk, is pressed against the mask from above using a linear slide and actuator.Fig. 5. Patterned mask on anode.Mask. As shown in Fig. 5, the mask consists ofa thin layer ofmaterial patterned with the desired features on a flat metal disk.

5. Patterned mask on anode.Mask. As shown in Fig. 5, the mask consists ofa thin layer ofmaterial patterned with the desired features on a flat metal disk. This particular mask incorporates several resolution test patterns, in which the smallest bar is 12 f..Ull wide and the largest is 200 f..Ull wide.Experiments were perfonned to detennine suitable masking conditions with the goal ofeliminating"flash" (extraneous deposits).Second-generation machine. We constructed the machine shown schematically in Fig. 6, to allow fabrication ofmulti-layer, 3-D parts. The machine consists ofseveral subsystems: the first includes a metal substrate onto whichthe layers are deposited, a linear slide to move the substrate up and down, an actuator, and an indicator to monitor lapping progress. Another subsystem includes a patterned anode (with multiple masks, one ofwhich is shown in Fig. 7), precision X and Y stages, and a tank to contain the electrolyte.

Another subsystem includes a patterned anode (with multiple masks, one ofwhich is shown in Fig. 7), precision X and Y stages, and a tank to contain the electrolyte. A third subsystem includes an anode and an electrolyte tank.165Z indicalorLapping plateXslage(a)(b)(c)Fig. 6. Schematic ofsecond-generation EFAB machine.The steps required to make a part are as follows:1. Lap the substrate to establish a reference/starting plane. 2. Position the anode to place the mask for the current layer below the substrate (Fig. 6(a)), then mate the substrate with the mask and selectively deposit several Jlm ofthe fIrst material.3: Deposit several J.lm of the second material over the fIrst material (Fig. 6(b)). 4. Lap the deposited layer (Fig. 6(c)) until the fInal desired thickness is reached. 5. Repeat steps 2-4 for remaining layers of the part. 6.Immerse the deposited material in an etchant to remove the sacrifIcial material.eu etching.

5. Repeat steps 2-4 for remaining layers of the part. 6.Immerse the deposited material in an etchant to remove the sacrifIcial material.eu etching. The ability to remove sacrificial material from narrow and long channels is a potential limitation on part geometry. We artificially created such geometry in advance ofhaving EFAB'd test parts by embedding Cu wires ofvarious diameters in epoxy with only their ends exposed, and immersing these in a commercial Cu etchant. We also immersed Ni wire in the etchant to see ifthere was any significant effect on Ni.ResultsInstant Masking. Using the first.,generation machine, we were able to obtain excellent defmition ofthepIated deposit: the edges of the deposit were sharp and there was virtually no flash. This is shown in Fig. 9 (traces ofelectrolyte remain on the substrate in this picture). Rectangular bars down to about 25 J.lffi width were well defined, although 12-J.lffi bars showed some definition problems.Fig. 7.

9 (traces ofelectrolyte remain on the substrate in this picture). Rectangular bars down to about 25 J.lffi width were well defined, although 12-J.lffi bars showed some definition problems.Fig. 7. Mask for selectively depositing sacrifIcial material on a part layer: the circle is 1.3 mm diameter.166Cu etching. We observed a reasonable etch rate ofthe epoxy-embedded Cu wires (200-300 ~), andfound that this varied little over a wire diameter range of40-200 IJID. No significant size change ofthe Ni wire was observed.Other perfonnance issues. We have not as yet evaluated layer thickness accuracy, mask registration accuracy, residual stress, or material.properties (though we expect these to be similar to commercialelectroformedNi).EFAD applicationsBroadly speaking, EFAB can be used to fabricatemicroscopic parts, microelectromechanical systems (MEMS), and mesoscopic parts and machines.

There is a large and growing market for such devicesin·such industries as medica4 automotive, electronics, and aerospace. Some application examples are:Micro- and mesoscopic mechanical machine parts, such as hinges, springs, screws, drive chains, and helical gears;Monolithically-manufactured systems such as motors, gearboxes, gas turbines, chemical processors, microsurgical and endoscopic instruments, optical scanners, motion sensors, compressors, pumps, cooling systems, solenoids, and valves; Fig. 8. Deposit produced using Instant Masking using first-generation machine and the mask shown in Fig. 5.Micro- and mesoscopic electrical parts, such as on-chip inductors, transformers, transmission lines, and heat sinks; electronic packaging components, such as interconnects for multichip modules.In our current program, we plan to use EFAB to produce a functional mesoscopic vibration sensor,shown as a CAD model in Fig. 9.Fig. 9.

9.Fig. 9. Preliminary concept of a vibration sensor (overall size 1 x 1 x 0.5 mm)167EFAB limitationsWhile well suited for small devices, electrodeposition is too slow (typically 25-100 prn/hr) a deposition process in most casesforthe productionofma.cr0scopicdevices(e.g.,larg~ than 1Omm in height). Fig. 10 S\lgg~.(lsui1llblesW;.range.forEF}\J3devi~s'ifurth~o~,thet'e isditlicuJty usin.S••~lectrodepositionto depos~materials oflow conductivity. Also, EFABalso requkespalt-specific tooling (photomasks), th~by addingtocost and.not allowingdevicefabrication to begin instantly after CAD. Finally, device geometries will notbeunlimited: removal ofsacrificial material will impose some constraints.ConclusionsWe aredevelopin.g a newSFF process and machinery based on selective elec1rod~sition. We have demonstratedthatwecan selectively deposit features down to about 25 Jllll in width with good defini,tion, andwea.reworkingto producemulti-layer parts.

We have demonstratedthatwecan selectively deposit features down to about 25 Jllll in width with good defini,tion, andwea.reworkingto producemulti-layer parts. In developing this process, we hope to provide a capable, genericmicrofabrication processthatwill allow fimctional metalmicro- and mesoscopic parts and complex systemsitobePrptQtypedaudmass-pro<l\lced by a fixed process that is cost-effective and speedy.

FECT OF A DISTRIBUTED HEAT SOURCE ON MELT POOL GEOMETRY AND MICROSTRUCTURE IN BEAM-BASED SOLID FREEFORM FABRICATIONSrikanth Bontha and Nathan W. Klingbeil Department of Mechanical and Materials Engineering Wright State University Dayton, OH 45435AbstractThe ability to control geometric and mechanical properties of parts fabricated using laser-based manufacturing processes requires an understanding and control of melt pool geometry and mi- crostructure. With the development of electron beam manufacturing or future beam-based depo- sition processes, the user may have more control over the distribution of incident energy, so that beam width becomes a potential process variable. As such, the focus of this work is the effect of a distributed heat source on melt pool geometry (length and depth) and the thermal conditions controlling microstructure (cooling rates and thermal gradients) in beam-based solid freeform fab- rication.

Previous work by the authors has employed the Rosenthal solution for a moving point heat source to determine the effects of process variables (laser power and velocity) on solidiﬁca- tion cooling rates and thermal gradients controlling microstructure (grain size and morphology) in laser-deposited materials. Through numerical superposition of the Rosenthal solution, the current work extends the approach to include the effects of a distributed heat source for both 2-D thin- wall and bulky 3-D geometries. Results suggest that intentional variations in beam width could potentially enable signiﬁcant changes in melt pool geometry without affecting microstructure.IntroductionOver roughly the past decade, a number of laser-based material deposition processes have been developed for automatic fabrication of complex parts, adding features to existing parts, or repair of worn components.

However, the widespread commercialization of these processes will ultimately require the ability to predict and control melt pool size, residual stress and microstructure through changes in process variables (e.g., laser power and velocity) [1]. With the advent of electron beam manufacturing or other next-generation processes, designers may have increased control over the distribution of incident energy compared to laser-based manufacturing. As a result, the ability to change the distribution of power (e.g., through changes in beam width) represents an additional process variable.Among the key deposit characteristics, the control of melt pool size assumes the highest pri- ority within the manufacturing community. This is because a consistent melt pool size is needed before speciﬁc features can even be built.

This is because a consistent melt pool size is needed before speciﬁc features can even be built. At the same time, the control of microstructure is also critical, particularly in aerospace and other structural applications that have strict guidelines on re- sulting mechanical properties. While the control of melt pool size and solidiﬁcation microstructure have been addressed in the literature [2], their interconnection has yet to be fully investigated.

While the control of melt pool size and solidiﬁcation microstructure have been addressed in the literature [2], their interconnection has yet to be fully investigated. In particular, it has not been shown how changing process variables to control melt pool size might simultaneously affect cooling rates and thermal gradients which ultimately control microstructure.To this end, the current study investigates how changes in laser or electron beam width simul- taneously affect melt pool geometry and microstructure in beam-based solid freeform fabrication.478The approach is based on superposition of the well known Rosenthal solution for a moving point heat source traversing an inﬁnite substrate [3], which has found extensive application in the lit- erature. While the Rosenthal solution was initially applied to welding, its application to laser deposition was ﬁrst carried out by Dykhuizen and Dobranich [4]. Vasinonta et al.

While the Rosenthal solution was initially applied to welding, its application to laser deposition was ﬁrst carried out by Dykhuizen and Dobranich [4]. Vasinonta et al. subsequently used the Rosenthal solution to guide the development of process maps relating laser deposition process variables to melt pool size and residual stress [5, 6]. The present authors have adopted a similar approach to the development of thermal process maps for predicting solidiﬁcation mi- crostructure in laser deposited materials [7–9]. While the authors’ prior work has been limited to a point heat source, the current study extends the approach to include the effect of a distributed heat source on both melt pool geometry and the thermal conditions controlling microstructure.

Results suggest that intentional variations in beam width could effect signiﬁcant changes in melt pool size without affecting microstructure.Geometries ConsideredThis study considers both the thin-wall (2-D) and bulky (3-D) geometries of Fig. 1, in which the process variables of interest are the absorbed power αQ and velocity V . In each case, it is assumed that the height h and length L are sufﬁciently large such that the steady-state Rosenthal solution for a moving point heat source applies [3].αQVαQVzoxozoxo yohhLLbb(a)(b)Figure 1: (a) Thin-Wall and (b) Bulky (3-D) Geometries ConsideredFormulation for 2-D Thin Wall GeometryThis section presents the formulation for a uniform power distribution of ﬁnite width w by superposition of the 2-D Rosenthal solution for a moving point heat source. The 2-D Rosenthal solution for the quasi-steady-state heat conduction problem of Fig. 1 (a) has been given in dimen- sionless form by Vasinonta et al.

The 2-D Rosenthal solution for the quasi-steady-state heat conduction problem of Fig. 1 (a) has been given in dimen- sionless form by Vasinonta et al. [5, 6] as (cid:2)(cid:3)(cid:4)T = e−x0K00 + z2 x2 0,where K0 is the modiﬁed Bessel function of the second kind, order zero. The dimensionless variables in eq. (1) are deﬁned in terms of the absorbed laser power αQ and velocity V asT =T − T0 αQ πkb,x0 =x0 2k ρcVand z0 =z0 2k ρcV.479(1)(2)In the above normalizations, T is the temperature at a location (x0,z0) relative to the moving point heat source and T0 is the initial temperature of the wall. The relative coordinates (x0,z0) are related to ﬁxed spatial coordinates (x, z) at any time t as (x0,z0)=(x−V t,z), where V is the laser velocity. Also in eq.

The relative coordinates (x0,z0) are related to ﬁxed spatial coordinates (x, z) at any time t as (x0,z0)=(x−V t,z), where V is the laser velocity. Also in eq. (2), the thermophysical properties ρ, c and k are the density, speciﬁc heat and thermal conductivity of the material, which are assumed to be temperature-independent.When the beam is modeled as a distributed heat source, the absorbed power αQ is distributed over a dimensionless width w as shown in Fig. 2 (a). The resulting distributed heat source is q = αQ/w. In keeping with the spatial normalizations of eq. (2), the normalized beam width is deﬁned as w = w/(2k/ρcV ).Q q (cid:68) (cid:32) wssdsq )((cid:32)Q (cid:68) wsd(cid:16)w 2w 2x0(cid:16)w 2w 20xz0z0(a)(b)Figure 2: Uniformly Distributed Heat Source for 2-D Thin Wall GeometryAs illustrated in Fig. 2 (b), the quantity q(s) ds acts as a point heat source at a distance s from the z0 axis.

2 (b), the quantity q(s) ds acts as a point heat source at a distance s from the z0 axis. Hence, the solution corresponding to each point heat source q(s) ds can be obtained by replacing x0with (x0 − s) in the dimensionless 2-D Rosenthal solution of eq. (1). The total solution is obtained by summing all the point heat sources q(s) ds between − w2 and w2 , i.e.,T =1 w(cid:5) w 2− w 2e−(x0−s) K0(cid:2)(cid:3)(x0 − s)2 + z2 0(cid:4)d¯s .Equation (3) represents the dimensionless 2-D Rosenthal solution for an absorbed power αQ uni- formly distributed over a beam width w.Representative Results for 2-D Thin Wall GeometryThis section includes representative dimensionless results which illustrate the effect of beam width on melt pool geometry and microstructure for 2-D thin wall geometries. For the purpose of illustration, attention is restricted to small-scale deposition of Ti-6Al-4V, although more compre- hensive results are given in [10].

For the purpose of illustration, attention is restricted to small-scale deposition of Ti-6Al-4V, although more compre- hensive results are given in [10]. The values of laser power (Q = 350 W) and velocity (V = 4.23 mm/s) considered here fall in the range of powers and velocities that are typical of those used in the LENSTM deposition of thin-wall geometries. The thermophysical properties of Ti-6Al-4V are assumed constant at the melting temperature Tm = 1654◦C, while the absorption coefﬁcient is as- sumed to be α = 0.35. Finally, the wall thickness is assumed constant at b = 2.26 mm, which is in keeping with that used in the authors’ prior published work [7–9,11].

Finally, the wall thickness is assumed constant at b = 2.26 mm, which is in keeping with that used in the authors’ prior published work [7–9,11]. The above values correspond to a dimensionless melting temperature T m = 2.88, whereT m =Tm − T0 αQ πkb.480(3)(4)Effect of Beam Width on Melt Pool GeometryIn order to determine the melt pool geometry (length and depth), the coordinates (x0, z0) which lie on the boundary of the melt pool are determined by replacing T with the melting temperature T m and ﬁnding the roots of eq. (3) numerically. In so doing, both the numerical integration and root ﬁnding are conducted using the software package MATLAB. The resulting normalized melt pool length ¯l and normalized melt pool depth ¯d are deﬁned asl =l 2k ρcVand d =d 2k ρcV,which are in keeping with the spatial normalizations of eq. (2).The normalized melt pool length and depth are plotted as a function of normalized beam width in Fig. 3.

3 (a) indicate that the normalized melt pool length initially increases as the laser power is changed from a point source (w = 0) to a uniformly distributed source of width w. This is because spreading out the heat source initially heats more material along the length to temperatures above T m. This increase in melt pool length continues until it reaches a maximum value at a speciﬁc value of w. At this point, the heat ﬂux becomes insufﬁcient to melt any additional material along the length, so that subsequent increases in w result in a decrease in melt pool length. This decrease in melt pool length continues until the heat ﬂux is no longer sufﬁcient to melt any material at all, so that the melt pool ceases to exist (l → 0).In contrast to the normalized melt pool length, the normalized melt pool depth decreases mono- tonically with increasing w (Fig. 3 (b)).

3 (b)). While spreading out the heat source initially melts more4810.35(5)0.4material in the length direction, the amount of heat available to melt the material through the depth decreases. As observed for melt pool length, this decrease in melt pool depth continues until the point when the melt pool ceases to exist (d → 0). Inspection of the ordinate scales in Fig. 3 suggests that melt pool depth is more sensitive to changes in beam width compared to melt pool length; however, results also suggest that changing the beam width can have a signiﬁcant effect on melt pool length, melt pool depth and the overall shape of the melt pool.Effect of Beam Width on MicrostructureAs discussed in [7–9], the microstructure (grain size and morphology) of deposited metals de- pends on cooling rates and thermal gradients at the onset of solidiﬁcation (i.e., along the boundary of the melt pool).

4 (a) reveal that increasing w results in an increase in both thermal gradient and cooling rate up to a certain depth within the melt pool, after which cooling rates and thermal gradients begin to decrease (relative to those for w = 0). This effect can be explained as follows. For the case of a point source (w = 0), the left hand boundary of the melt pool (i.e., the solidi- ﬁcation boundary) is always relatively far from the heat source. As the laser power is distributed (increasing w), the solidiﬁcation front near the surface begins to see the edge of the distributed source, which leads to an increase in both cooling rate and thermal gradient. However, beyond a certain depth within the melt pool (i.e., further away from the edge of the distributed source), bulk heating of the melt pool due to the distributed source results in a decrease in both cooling rate and thermal gradient compared to their point source counterparts.

Given the solidiﬁcation cooling rate ∂T ∂t and thermal gradient G = |∇T|, the solidiﬁcation rate R is determined asR =1 G∂T ∂t.483(8)The expected grain morphology can be predicted as either equiaxed, columnar or mixed by plotting points in G vs. R space (i.e., on the “solidiﬁcation map”), which has been previously calibrated for Ti-6Al-4V [14].A solidiﬁcation map showing the effect of beam width on predicted grain morphology for small-scale (low power) deposition of thin-wall Ti-6Al-4V deposits is shown in Fig. 5. The values of G and R plotted in Fig. 5 are extracted from the dimensionless cooling rate and thermal gradient plots of Fig. 4, with thermophysical properties at 1654◦C and process variables as previously noted.From Fig. 5, it is clear that the data points corresponding to all beam widths fall in the fully columnar region.

5, it is clear that the data points corresponding to all beam widths fall in the fully columnar region. This result is in keeping with the solidiﬁcation map predictions previously re- ported for a point heat source [7–9, 15], as well as with experimental observations for LENSTM deposited Ti-6Al-4V [15,16]. More importantly, the results of Fig. 5 reveal that in contrast to melt pool geometry, solidiﬁcation grain morphology is insensitive to changes in beam width.Formulation for Bulky 3-D GeometryThis section presents the formulation for a uniform distribution of power over a circular beam of ﬁnite width w by superposition of the 3-D Rosenthal solution for a moving point heat source. The formulation for a square beam can be found in [10].0x0y0zr0xdrdrqy00z(a)(b)Figure 6: Uniformly Distributed Heat Source for Bulky 3-D Geometry (Circular Beam Proﬁle)The 3-D Rosenthal solution for the quasi-steady-state heat conduction problem of Fig.

1 (b) isgiven in dimensionless form as“√”0+y2 x20+z2 0e− 2x0+T =(cid:7).x2 0 + y2 The dimensionless variables in eq. (9) are deﬁned in terms of the absorbed laser power αQ and velocity V as0 + z2 0T =T − T0 )( ρcV 2k( αQ πk),x0 =x0 2k ρcV,y0 =y0 2k ρcVandz0 =z0 2k ρcV.The above deﬁnition of T is in keeping with that used for bulky 3-D geometries in [5, 7, 13, 17], and differs by a factor of 2 from a more recent deﬁnition used by other researchers [18–20].484(cid:32)(cid:84)qdA(9)(10)When the beam is modeled as a uniform distributed heat source with a circular beam proﬁle, the absorbed power αQ is distributed over a circle of dimensionless diameter w as shown in Fig. 6 (a). The resulting distributed heat source is q = αQ/(πw2/4). As illustrated in Fig. 6 (b), the quantity q r dr dθ acts as a point heat source at a location (r, θ) from the origin.

6 (a). The resulting distributed heat source is q = αQ/(πw2/4). As illustrated in Fig. 6 (b), the quantity q r dr dθ acts as a point heat source at a location (r, θ) from the origin. Hence, the solution corresponding to each point source q rdr dθ can be obtained by substituting αQ = q rdr dθ, x0 = x0 − r cos θ into the dimensionless 3-D Rosenthal solution of eq. (9).

Attention is largely restricted to small-scale deposition of Ti-6Al-4V, while more general results are provided in [10]. The values of laser power (Q = 550 W) and velocity (V = 8.47 mm/s) considered here fall in the range of powers and velocities that are typical of those used in LENSTM deposition of bulky 3-D geometries.

7 correspond to actual beam widths in the range 0 < w < 2.0 mm. While trends in the normalized melt pool length and depth are similar to their 2-D counterparts of Fig. 3, the melt pool geometry for bulky 3-D geometries is somewhat less sensitive to variations in beam width. Still, it is clear that variations in beam width can effect signiﬁcant changes in melt pool geometry for bulky 3-D deposits.Effect of Beam Width on MicrostructureThe expressions for the dimensionless cooling rate and thermal gradient can be obtained through differentiation of eq. (11), and are given in [10]. Representative results for the normalized cool- ing rate and thermal gradient on the solidiﬁcation boundary are plotted in Fig. 8 as a function of relative depth within the melt pool for selected values of the beam diameter w.

9 (a) (small-scale process), it is clear that the data points corresponding to all beam widths and depths fall in the fully columnar region. This result is in keeping with the solidiﬁcation map predictions previously reported for a point heat source [7, 8], as well as with experimental observations for LENSTMdeposited Ti-6Al-4V [14, 21]. The results of Fig. 9 (b) (large-scale process) suggest the possibility of a graded microstructure through the depth of the deposit, with a mixed or even fully equiaxed microstructure at the surface. This result is again in keeping with the solidiﬁcation map predictions previously reported for a point heat source [7, 8], as well as with experimental observations reported in the literature for a 14 kW large-scale process [14]. While spreading out the heat source (increasing w) moves the data slightly toward the fully columnar regime, the results of Fig.

While spreading out the heat source (increasing w) moves the data slightly toward the fully columnar regime, the results of Fig. 9 appear largely insensitive to variations in beam width.Summary and ConclusionsThe ability to produce parts with required geometric and mechanical properties using beam- based solid freeform fabrication requires an understanding and control of melt pool geometry and microstructure. With the development of electron beam manufacturing or other next-generation beam-based deposition processes, the user may have more precise control over the distribution of incident energy (e.g., beam width). Through superposition of the Rosenthal solution for a mov- ing point heat source, the current study has investigated the effect of beam width on melt pool48710geometry (length and depth) and the thermal conditions controlling microstructure (grain size and morphology) for both 2-D thin wall and bulky 3-D geometries.

The results of this work suggest that variations in beam width can effect signiﬁcant changes in melt pool geometry without altering the microstructure.AcknowledgmentsThis work has been supported by the National Science Foundation, grant number DMI-0224517, by the joint AFRL/DAGSI Research Program, project number ML-WSU-01-11, and by Wright State University and the Ohio Board of Regents.References[1] J. Beuth and N. Klingbeil, “The Role of Process Variables in Laser-Based Direct Metal SolidFreeform Fabrication,” JOM, pp. 36–39, September 2001.[2] N. Klingbeil, S. Bontha, D. Gaddam, C. Brown, J. Beuth, A. Birnbaum, and P. Aggarangsi, “Modeling of Melt Pool Size and Solidiﬁcation Microstructure in Laser-Based Additive Man- ufacturing,” in Proceedings 2006 NSF DMII Grantees and Research Conference, (St. Louis, MO), July 2006.[3] D. Rosenthal, “The Theory of Moving Sources of Heat and Its Applications to Metal Treat-ments,” Transactions of ASME, vol. 68, pp.

Fraser, “Prediction of Microstructure in Laser Deposition of Titanium Alloys,” in Solid Freeform Fabrication Proceedings, (Austin, TX), pp. 142–149, August 2002.[12] A. Vasinonta, J. Beuth, and M. Grifﬁth, “Process Maps for Laser Deposition of Thin-Walled Structures,” in Solid Freeform Fabrication Proceedings, (Austin, TX), pp. 383–391, August 1999.[13] A. Vasinonta, Process Maps for Melt Pool Size and Residual Stress in Laser-Based SolidFreeform Fabrication. PhD thesis, Carnegie Mellon University, May 2002.[14] P. Kobryn and S. Semiatin, “Microstructure and Texture Evolution during Solidiﬁcation Pro- cessing of Ti-6Al-4V,” Journal of Materials Processing Technology, vol. 135, pp. 330–339, 2003.[15] C. J. Brown, “Modeling of Solidiﬁcation Microstructure in Laser-Deposited Ti-6Al-4V,”Master’s thesis, Wright State University, June 2003.[16] X. Wu, J. Liang, J. Mei, C. Mitchell, P. Goodwin, and W.

J. Brown, “Modeling of Solidiﬁcation Microstructure in Laser-Deposited Ti-6Al-4V,”Master’s thesis, Wright State University, June 2003.[16] X. Wu, J. Liang, J. Mei, C. Mitchell, P. Goodwin, and W. Voice, “Microstructures of Laser-Deposited Ti-6Al-4V,” Materials and Design, vol. 25A, pp. 137–144, 2004.[17] A. Vasinonta, J. Beuth, and R. Ong, “Melt Pool Size Control in Thin-Walled and Bulky Parts via Process Maps,” in Solid Freeform Fabrication Proceedings, (Austin, TX), pp. 432–440, August 2001.[18] A. Birnbaum, J. Beuth, and J. Sears, “Scaling Effects in Laser-Based Additive ManufacturingProcesses,” in Solid Freeform Fabrication Proceedings, (Austin, TX), August 2004.[19] A. Birnbaum, P. Aggarangsi, and J. Beuth, “Process Scaling and Transient Melt Pool Size Control in Laser-Based Additive Manufacturing Processes,” in Solid Freeform Fabrication Proceedings, (Austin, TX), pp. 328–340, August 2004.[20] P. Aggarangsi, J. Beuth, and D.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceTHE EFFECT OF ARC-BASED DIRECT METAL ENERGY DEPOSITION ON PBF MARAGING STEEL Bishal Silwal*, Christopher Gerdmann*, Stephen Migues*, Kamram Kardel**, Shaowen Xu* Andrew Durrence***Department of Mechanical Engineering, Georgia Southern University, Statesboro, GA, 30460 **Department of Manufacturing Engineering, Georgia Southern University, Statesboro, GA, 30460 *** Gulfstream Aerospace, Savannah, GA AbstractWith the growing demand of metal based additive manufacturing it is inevitable that the additively manufactured parts will be used in consolidation with different manufacturing process. Arc based additive manufacturing is a process used to produce three-dimensional structure using welding arc as a power source.

Arc based additive manufacturing is a process used to produce three-dimensional structure using welding arc as a power source. Powder Bed Fusion (PBF) is an additive manufacturing process using laser power to generate a three dimensional structure. In this paper, the effect of direct energy deposition on the PBF has been investigated. The microstructure changes and the melting region are characterized. A computational fluid dynamics model is used to predict the melt region and temperature in PBF.Keywords: Additive manufacturing, welding arc, GTAW, ABAM, Maraging Steel, PBFIntroductionThe additive manufacturing (a.k.a 3d printing) industry generated over $5 billion in 2015 and is expected to grow over five times by year 2021 [4]. According to ASTM, there are seven categories of additive manufacturing: photopolymerization, material jetting, binder jetting, material extrusion, powder bed fusion, sheet lamination and direct enegy deposition.

Based upon the power source and feed material, it can be further classified into many individual processes (Figure 1).Figure 1: Additive manufacturing technology based upon the power source and feed material shape [12]1585Powder bed fusion (PBF) is one of the many commercially available additive manufacturing (a.k.a 3-d printing) technologies to fabricate net shape parts. In this process, a thin layer of powder on powder-bed is slectively melted by laser, building a volume on a layer-by-layer basis. The major advantage of using PBF is that a complex geometry can be acheived which may yield longer time by using other conventional manufacturing technologies.Arc based additive manufacturing (ABAM) is a process in which a welding arc used as a power source to melt the substrate and filler-wire. The tool path can be generated via an industrial robot, numerical controlled gantry system or a Rep-Rap system.

The tool path can be generated via an industrial robot, numerical controlled gantry system or a Rep-Rap system. The deposition rate of 10-20 lbs/hr can be achieved in this process, which is promising in large structure fabrication. Two of the main challenges restricts its’ use in AM industry are lower resolution of layer height (compared to PBF) and lack of technical resource (mainly path planning) for complex geometry.Maraging steels offer combination of excellent material properties such as tensile strength, fracture toughness, dimensional stability and good weldability. These steels are extensively used in aerospace and tooling industries. There are four principal grades of wrought 18% nickel maraging steels; designated as Grades 200, 250, 300 and 350 based upon their tensile strength value in English units. The primary microstructure of the maraging steel is a martensite matrix with precipitation of intermetallic compounds such as Ni3Ti, Fe2Mo etc.

The primary microstructure of the maraging steel is a martensite matrix with precipitation of intermetallic compounds such as Ni3Ti, Fe2Mo etc. It has also reported that due to formation of retained austenite during aging, the mechanical properties such as tensile and toughness improves [5]. Many researchers have investigated the mechanical properties of the DMLS produced maraging steel [1, 3]. The precipitation hardening behavior has also been studied in the past [2]. Based upon the observation from [2], the as-fabricated PBF maraging steel has martensite matrix with small regions of retained austenite. In order to improve the structural properties, further processing such as shot peening, welding and heat treatment etc of PBF fabricated part are unavoidable in manufacturing. In this paper, the effect of welding arc deposited maraging wire on PBF fabricated part is studied. A cold-wire Gas Tungsten Arc Welding (GTAW) has been used as a heat source.

In this paper, the effect of welding arc deposited maraging wire on PBF fabricated part is studied. A cold-wire Gas Tungsten Arc Welding (GTAW) has been used as a heat source. The surface texture, macro and microstructure and hardness of PBF fabricated maraging steel have been carried out.The numerical modeling in PBF is less computationally efficient due to the nature of the complexity and fineness involved. Lee et al. [6] has used the mesoscale model to predict the fluid flow, temperature and solidification in nickel based alloy. The powder packing model was developed using the discrete element method (DEM) [6]. More detail literature review on the aspect of numerical modeling has been found on [7, 8, and 9]. In this paper, a computationally efficient model is developed using a uniform particle bed, recoil pressure, evaporative pressure, Fresnel reflection and Gaussian distribution of laser heat source.ExperimentalThe experiment was performed in EOSINT M 280.

The material composition of maraging part (Grade 300) is presented in Table 1. The EOS parameters are presented in Table 2. The built volume was 4.00×1.75×0.30 inches. The surface texture was measured by using Keyence digital profilometer. After built, one part was cut off (0.50×1.75×0.30) from the base metal (substrate) using a horizontal band saw. The layer deposition was performed using cold-wire GTA on the part with the maraging wire identical in composition to maraging powder. The parameters for GTA process were: 3 mm/sec travel speed, 150 A welding current and 26 mm/sec wire feed speed.

The parameters for GTA process were: 3 mm/sec travel speed, 150 A welding current and 26 mm/sec wire feed speed. An argon gas with flow rate of 15 liter per minute was used as a shielding gas.1586Table 1: The chemical composition of the Grade 300 maraging steelGrade Maraging 0.05- 0.15AlC 0.003 8.5- 9.5Co Fe Mn Mo NiBal 0.14.60 5.2017.0- 19.0P Ti 0.01 0.01 0.1 0.6- 0.8SSiTable 2: The parameters in DMLSValues Parameters Powder size ~30 micron Focus Diameter 100-500 Ωm Laser power Scan speed GTA travel speed GTA Power Input GTA Wire feed speed200 W NA 3 mm/sec2700 W26 mm/secResults and DiscussionThe as-built unpolished surface profile measurement is presented in Figure 2. The surface roughness has the maximum value of 25.73 µm and minimum value of 20.38 µm. The micrograph is presented in the right side of the Figure 2, while the contour variation of surface topography is presented in the left side of the Figure 2.

The micrograph is presented in the right side of the Figure 2, while the contour variation of surface topography is presented in the left side of the Figure 2. The arrow indicates the built direction. The darker regionsFigure 2: The surface topography measurement of PBF maraging steel1587as seen on the micrograph have higher value while the brighter regions have lower value. It may be due to re-melting of the layer, which let the fluid to flow outward. The average layer thickness is ~ 100 micron. The adsorbed powders are also visible on the micrograph (Figure 3)Figure 3: The built direction with multiple layers on left and higher magnification on the right depicting the non- uniformity of the individual layer. It is also clear from the figure that the layers are non-uniform. The primary reason is due to powder size distribution and powder arrangement which affect the laser power absorption.

It is also clear from the figure that the layers are non-uniform. The primary reason is due to powder size distribution and powder arrangement which affect the laser power absorption. In addition, melting of powder, remelting of the solidified layer and fluid flow of the molten metal due to Maragoni effect and capillary effect play important role in the final shape of the solidified layer.Figure 3: The built direction with multiple layers on left and higher magnification on the right depicting the non- uniformity of the individual layer and adsorbed powderThe Rockwell hardness is presented in Figure 3. The hardness value is within the range of 31 to 37 HRC in the PBF maraging steel. The Gas Tungsten Arc (GTA) deposited maraging steel has the hardness value ranges from 16 to 23 HRC. The heat affected region of maraging steel has hardness of 28-29 HRC.

The Gas Tungsten Arc (GTA) deposited maraging steel has the hardness value ranges from 16 to 23 HRC. The heat affected region of maraging steel has hardness of 28-29 HRC. The GTA heat has definitely decrease the hardness of the PBF maraging steel from 34 average HRC to 28 HRC.Figure 4: Rockwell hardness measurement: Powder Bed Fusion (PBF) maraging steel on left and GTA direct deposition on the right (the darker area represents the layer deposited by GTA)1588Figure 5: The micrograph of the GTA deposited maraging steelThe GTA deposited surface is presented in Figure 5. The layer is nearly 800 µm, which is 8-9 times bigger than the average layer of PBF. A metallurgical bonding has been observed. The solidification morphology entails an epitaxial growth with planar-columnar in the joint, then columnar and equiaxed dendrite towards the outward region as depicted in Figure 5.

The solidification morphology entails an epitaxial growth with planar-columnar in the joint, then columnar and equiaxed dendrite towards the outward region as depicted in Figure 5. The solidification morphology of the PBF maraging steel is not conclusive, however some columnar grain is visible in the micrograph presented in the Figure 6. The energy-dispersive X-ray spectroscopy (EDX) analysis was performed on the PBF layers. The results, mainly chemical composition is presented in Table 3. There is no significant difference, however small difference of nickel content was observed.

The results, mainly chemical composition is presented in Table 3. There is no significant difference, however small difference of nickel content was observed. The area ‘A’ represents some fine grains (not visible in Figure 6) while ‘B’ represents columnar grains.Figure 6: The micrograph of PBF maraging steel, wire EDM cut, polished and etched (left), transverse direction on the right at higher magnification1589ABTable 3: EDX analysis of the PBF layersMap Sum SpectrumA (Wt%)B (Wt%)O1.161.56F1.18NAAl0.370.22Si0.280.28Ti0.620.58Fe66.8167.44Co7.087.28Ni15.3414.65Mo3.973.91SNA0.05Mathematical ModelThe mathematical model for GTA deposition has been presented in [11]. In this paper, the phenomenon in PBF with high heating of laser power has been considered. The laser sintering involves laser-powder interaction, phase change, heat and mass transfer and fluid flow.

In this paper, the phenomenon in PBF with high heating of laser power has been considered. The laser sintering involves laser-powder interaction, phase change, heat and mass transfer and fluid flow. A computation domain of 1000 х 140 х 120 micron is considered. A 30 micron layer of uniform diameter powder particles is considered. The laser beam profile is assumed as a Gaussian type axisymmetric distribution. The distribution equation of laser energy is described as𝑞(𝑟) = 𝑄𝑒(−𝑟 𝑟𝑏)2where, q denotes heat density function, Q is the total power of the laser beam, r is the coordinate in radial direction and rb is the laser beam radius at focus plane. The melting temperature and melt pool velocity are calculated using the conservation of momentum, conservation of energy and material continuity equations. The convection, conduction, radiation and evaporative boundary conditions are applied. The surface tension force is also applied as a boundary condition.

The convection, conduction, radiation and evaporative boundary conditions are applied. The surface tension force is also applied as a boundary condition. The VOF method is used to track the free surface. The numerical model consist of two region of 30 mm x 4 mm x 3 mm and constant diameter powder region with same stack size. The recoil pressure is calculated using,𝑃𝑟 = 𝐴𝑒{𝐵(1−𝑇𝑣 𝑇 )}Where, Pr is recoil pressure, Tv is evaporation temperature, T is the surface temperature,A and B are material dependent coefficients taken as 3500 and 5. The absorption rate (Fresnel reflection) is calculated from1590𝑎 = 1 −1 2(1 + (1 − 𝜀𝑐𝑜𝑠𝜙)2 1 + (1 + 𝜀𝑐𝑜𝑠𝜙)2 +𝜀2 − 2𝜀𝑐𝑜𝑠𝜙 + 2𝑐𝑜𝑠2𝜙 𝜀2 + 2𝜀𝑐𝑜𝑠𝜙 + 2𝑐𝑜𝑠2𝜙)where, ε is the coefficient dependent upon laser type, φ is the angle between incident and surface normal. The heat flux is further calculated using, 𝑞𝑛 = 𝑞𝑛−1 ∗ 𝛼𝑛.

The heat flux is further calculated using, 𝑞𝑛 = 𝑞𝑛−1 ∗ 𝛼𝑛. The loop is performed until the heat flux becomes small.,The saturated vapor pressure is given by,1 𝑇𝑏 Where, Ps is the saturated vapor pressure, P is the atmospheric pressure, Tb is the boiling1 𝑇)} −{𝜆(𝑃𝑠 = 𝑃𝑒temperature, T is the surface temperature, and 𝜆 is the evaporation energy per atom. The equation below is used to calculate the mass transfer due to evaporation and calculated based on the kinetic theory𝑀𝑛𝑒𝑡 = 𝑅𝑎𝑐𝑐𝑜𝑚√𝑀 2𝜋𝑅𝑇(𝑃𝑠 − 𝑃𝑣)Boundary Condition:𝐾𝜕𝑇 𝜕𝑛⃗+ ℎ(𝑇𝑠 − 𝑇) + 𝑄𝑣𝑎𝑝 + 𝜎𝜀(𝑇𝑠 𝐹 = 𝑛⃗ ∇(∇𝛾) − ∇𝑠𝛾4 − 𝑇4) = 𝑞(𝑟)The first boundary equation is for the energy equation, where K is the thermal conductivity and assume to be constant, n is the normal vector, Ts is the surface temperature, Qvap is heat loss due to vaporization, σ is the Boltzmann Constant, ε is the emissivity.

The flow is more dependent upon theMaragoni convection and vapor pressure due to initial evaporation temperature assumption of ~ 2300 K for maraging steel. An additional melt region may have been achieved due to Fresnel reflection.1 mmFigure 7: Temperature profile on the left and melt pool depth on the rightConclusionOne layer of maraging steel has been deposited on the PBF maraging steel by arc based additive manufacturing (ABAM). The microstructure of the layer is martensitic with some retained austenite. The epitaxial growth follows a planar, columnar and equiaxed grain size. The Rockwell hardness value has been reduced due to the heat of ABAM process, thus softening and removing the laser (fish-scale) tracks on the PBF maraging steel. A high fidelity numerical model has been developed to predict the temperature and the melt pool shape for the PBF process.

A high fidelity numerical model has been developed to predict the temperature and the melt pool shape for the PBF process. Future work will be focused on microstructural modeling and energy-dispersive X-ray spectroscopy analysis to differentiate the arc weld against the PBF maraging steel.AcknowledgementThe authors would like to thank Dr. Rafael Quirino and Ms. Khristal Monroe from Chemistry Department at Georgia Southern for providing assistance in SEM use.1592References1. Kempen, K., Yasa, E., Thijs, L., Kruth, J. P., & Van Humbeeck, J. (2011). Microstructure and mechanical properties of Selective Laser Melted 18Ni-300 steel. Physics Procedia, 12, 255-263.2. Jägle, E. A., Choi, P. P., Van Humbeeck, J., & Raabe, D. (2014). Precipitation and austenite reversion behavior of a maraging steel produced by selective laser melting. Journal of Materials Research, 29(17), 2072-2079.3. Tan, C., Zhou, K., Tong, X., Huang, Y., Li, J., Ma, W., & Kuang, T. (2016).

fect of Architecture and Porosity on Mechanical Properties of Borate Glass Scaffolds Made by Selective Laser SinteringKrishna C. R. Kolan1, Ming C. Leu1, Gregory E. Hilmas2 and Taylor Comte1, 31Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409 2Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409 3Department of Biomedical Engineering, University of North Carolina, Chapel Hill, NC 27514AbstractThe porosity and architecture of bone scaffolds, intended for use in bone repair or replacement, are two of the most important parameters in the field of bone tissue engineering. The two parameters not only affect the mechanical properties of the scaffolds but also aid in determining the amount of bone regeneration after implantation.

The two parameters not only affect the mechanical properties of the scaffolds but also aid in determining the amount of bone regeneration after implantation. Scaffolds with five different architectures and four porosity levels were fabricated using borate bioactive glass (13–93B3) using the selective laser sintering (SLS) process. The pore size of the scaffolds varied from 400 to 1300 µm. The compressive strength of the scaffolds varied from 1.7 to 15.5 MPa for porosities ranging from 60 to 30%, respectively, for the different architectures. Scaffolds were soaked in a simulated body fluid (SBF) for one week to measure the variation in mechanical properties. The formation of the Hydroxyapatite and in-vitro results are provided and discussed.1. IntroductionThe discovery of Bioglass® by Prof. Hench in 1969 has led to active research interest in the field of bioactive materials in the past four decades [1, 2].

IntroductionThe discovery of Bioglass® by Prof. Hench in 1969 has led to active research interest in the field of bioactive materials in the past four decades [1, 2]. The bioactive materials convert to Hydroxyapatite (HA), the main mineral constituent of bone, when exposed to body fluids thereby integrating with the surrounding tissue. Recently, interest has been focused on developing bioactive glasses as they offer excellent bioactive characteristics when compared to glass-ceramics or ceramics. Borate based bioactive glasses not only bond to the surrounding hard tissue but are also known to bond with soft tissues [3]. Table 1 shows the compositions (in wt.%) of the borate based 13-93B3 bioactive glass when compared to the silicate based 13-93 bioactive glass. The 45S5 glass composition is also presented for comparison. The molar concentration of SiO2 in 13-93 glass is replaced by B2O3 in the 13-93B3 glass.

The 45S5 glass composition is also presented for comparison. The molar concentration of SiO2 in 13-93 glass is replaced by B2O3 in the 13-93B3 glass. Borate glass is chemically less durable when compared to the silicate glass and therefore, it converts to HA at a faster rate and allows faster bone formation when compared to the silicate glass [3, 4].Table 1. Compositions (in wt.%) of 13-93 and 13-93B3 glasses compared to 45S5 glass. B2O3 - - 56.6SiO2 45 53 -P2O5 6 4 3.7CaO 24.5 20 18.5MgO - 5 4.6Na2O 24.5 6 5.5K2O - 12 11.145S5 13-93 13-93B3Recently, 13-93B3 glass scaffolds, with ~50% porosity, were fabricated with an organic based paste composition using the Robocasting technique [5, 6]. Organic solvents were used inthe above studies to prepare the paste because of the reactivity of borate glass with aqueous based solvents.

Organic solvents were used inthe above studies to prepare the paste because of the reactivity of borate glass with aqueous based solvents. Although porosity and pore size can be controlled using Robocasting, the process has limited control over the pore architecture when fabricating porous parts because of the layer- by-layer filament deposition. In comparison, powder based additive manufacturing (AM) techniques like the selective laser sintering (SLS) process provide flexibility in fabricating scaffolds with complex pore architectures as they do not require support structures during part fabrication. Therefore, the SLS process provides an opportunity to investigate the effects of porosity and pore architecture on the structural and biological characteristics of the scaffolds. There have been some articles in the literature wherein researchers have proposed techniques to develop the CAD models for scaffolds which closely mimic the human trabecular bone architecture [7-9].

There have been some articles in the literature wherein researchers have proposed techniques to develop the CAD models for scaffolds which closely mimic the human trabecular bone architecture [7-9]. However, fabricating scaffolds with such a complex architecture is still a challenge as not all AM techniques can fabricate them with bioceramics at the required resolution [10]. Also, limited work has been done to compare the bioactive scaffolds made with complex architectures and the traditional lattice structures, considering the aspects of both manufacturability and their behavior after in vitro or in vivo tests. In our previous work, we have shown that silicate based 13-93 bioactive glass scaffolds made by the SLS process provide good mechanical properties and preferable surface morphology for cell proliferation [11-13].

A low laser power and energy density was employed during the fabrication, resulting only melting the polymeric binder and assisting in fusion of the bioactive glass particles without affecting the material properties of the borate glass. Here, we hypothesized that scaffolds can be fabricated with 13-93B3 bioactive glass using the same processing parameters as established in our previous work [13].In the current work, we investigated the effects of pore architecture and porosity on the mechanical properties of the scaffolds. Five different architectures were considered and the scaffolds were fabricated with each of these architectures at four designed porosity levels. The sintered scaffolds were immersed in simulated body fluid (SBF) for one week and the effects of architecture and porosity on the compressive resistance of the scaffolds were studied. The materials and methods used for this study are noted in Section 2.

The materials and methods used for this study are noted in Section 2. The results obtained, including the variation of compressive strength with respect to porosity and architecture, before and after immersion in SBF, are presented in Section 3.2. Materials and Methods2.1. Fabrication of scaffolds13-93B3 bioactive glass (prepared by Mo-Sci Corp., Rolla, MO) with a nominal chemical composition of 56.6% B2O3, 3.7% P2O5, 18.5% CaO, 4.6% MgO, 5.5% Na2O and 11.1% K2O (by weight) was used in this research. The average particle size of the 13-93B3 particles was measured to be ~12 µm (d50), with a d10 of 4 µm and a d90 of 30 µm. Particle size distributions were obtained using a laser diffraction-based particle size analyzer (S3500, Microtrac Inc., Largo, FL).

Particle size distributions were obtained using a laser diffraction-based particle size analyzer (S3500, Microtrac Inc., Largo, FL). The 13-93B3 glass particles were mixed with stearic acid as the binder (C18H36O2, grade HS, Acros Organics, Morris Plains, NJ) and dry ball-milled for 8 hrs with ZrO2 milling media to obtain the feedstock powder for the SLS machine. A 15% binder content by weight was used for this purpose as this quantity provides the best possible results and is based on our previous study of SLS using 1393 bioactive glass [13]. The fabrication experiments were carried out on a commercial DTM Sinterstation 2000 machine. The effect of SLS parameters onfabricating scaffolds using stearic acid binder and bioglass powders has already been investigated in our previous work and the same set of parameters (laser power – 5 W, scan speed – 508 mm/s, scan spacing – 0.23 mm, layer thickness – 76.2 µm) were adopted for the current study [13].2.2.

Post-processing and physical assessment of scaffoldsThe fabricated green parts were post-processed in a three-stage programmable air furnace (Vulcan Benchtop Furnace, York, PA). The following heat treatment schedule was used for this study: de-binding heating rate of 0.1°C/min to 550°C; increase in the heating rate to 1°C/min until a final sintering temperature of 570°C with a 1 hour hold; and then the furnace was turned off and allowed to cool to room temperature. Optical microscopy was used to measure the pore sizes of the sintered scaffolds. Archimedes method was used to measure the apparent porosity of the sintered scaffolds. Cubic shaped parts measuring 1 cm in length were used for the purpose of measuring porosity and parts measuring 5 mm in length were used for the purpose of mechanical testing and SBF study. A cross-head speed of 0.5 mm/min was used on a mechanical load frame (Instron 4469 UTM, Norwood, MA) to determine the compressive strengths of the parts.

A cross-head speed of 0.5 mm/min was used on a mechanical load frame (Instron 4469 UTM, Norwood, MA) to determine the compressive strengths of the parts. Five samples in each set were used, and the results were reported as the average value ± the standard deviation. Scans were run from 2θ values ranging from 10o to 80o using Cu Kα radiation (λ = 0.154056 nm) for powder X-ray diffraction (XRD) analysis (Philips X-Pert, Westborough, MA) on the as-received 13-93B3 powder, sintered scaffolds and also on the dried scaffolds after removing them from the SBF to determine the changes in the crystalline/amorphous nature of the material.2.3. SBF testsThe SBF solution was prepared based on the Kokubo method [14]. All the samples were ultrasonically cleaned three times using ethanol and then dried in an oven overnight before being kept in the SBF solution. The scaffolds were weighed and 100 ml of solution was used for 1 gram of the scaffold for soaking.

The scaffolds were weighed and 100 ml of solution was used for 1 gram of the scaffold for soaking. The SBF solution container with scaffolds was kept in an incubator maintained at 37°C. All the compression tests were conducted on wet scaffolds to provide realistic data on the degradation of the scaffolds. Three samples in each set were used for compression testing and the results are reported as the average value ± the standard deviation. The scanning electron microscopy (SEM) (S-570, Hitachi Co., Tokyo, Japan) images were obtained to analyze the surface morphology of the scaffolds.3. Results and Discussion3.1. Fabrication of ScaffoldsThe CAD models of the repeatable units of the five architectures considered in this work are shown in Figure 1(a). The first unit cube is modeled by the solid tubular structures which are perpendicular to each other and running in x, y, and z directions.

The first unit cube is modeled by the solid tubular structures which are perpendicular to each other and running in x, y, and z directions. The porosity in this “cubic” architecture is a function of the diameter of the tube and distance between the tubes. The second unit cube is formed by subtracting a sphere from a solid cube and is termed as “spherical” pore architecture based on the geometry of the pore formed because of the design. The porosity in thiscase is a function of the diameter of the sphere and length of the unit cube. The third unit cube is formed by diagonally joining the corners of a unit cube with tubular structures and is termed the “X” architecture. The porosity of the “X” architecture is a function of diameter of the tubes. The latter three architectures are the typical and frequently used designs in most of the AM techniques to manufacture scaffolds.

The latter three architectures are the typical and frequently used designs in most of the AM techniques to manufacture scaffolds. The fourth and fifth are “diamond” and “gyroid” architectures, which are based on the freeform surface and try to mimic the trabecular architecture of bone. The surface was generated using open source software K3DSurf v0.6.2 [15]. The governing mathematical equations of these architectures can be modified to establish a range of porosities. The equation for gyroid surface is represented as:cossincossincossin0xyyzzxk, where, 0 ≤ k ≤ 1 and –π ≤ x, y, z ≤ π;and for a diamond surface, the governing equation is represented as:sinsinsinsincoscoscossincoscoscossin0xyzxyzxyzxyzk, where, 0 ≤ k ≤ 1 and0 ≤ x, y, z ≤ π;The surface generated was modeled into a volume with the help of pre-processor software and then converted to a .STL file for fabrication using the SLS machine.

The aspect of manufacturability was considered during the design process for all the scaffolds. For example, in the spherical architecture, a reduction in porosity was achieved by decreasing the diameter of the sphere to be subtracted from the cube. However, decreasing the diameter of the sphere will decrease the pore diameter and removing unsintered powder from fine pores after fabrication becomes limited, increasing the likelihood of damaging the scaffold structure. Therefore, the pore sizes of all the scaffolds were designed in such a way that the unsintered powder could be removed after the scaffold fabrication. The fabricated scaffolds after binder burnout and sintering are shown in Figure 1(b) as representative specimens from each of the architectures in the same order as shown in Figure 1(a).Figure 1.

The fabricated scaffolds after binder burnout and sintering are shown in Figure 1(b) as representative specimens from each of the architectures in the same order as shown in Figure 1(a).Figure 1. (a) CAD models of the repeatable units of five architectures: (i) cubic, (ii) spherical, (iii) X, (iv) diamond, and (v) gyroid; (b) 13-93B3 scaffolds with different architectures after sintering.The porosities of all the five architectures were designed to be 50, 60, 70, and 80%. However, as mentioned earlier, several factors like laser spot diameter, effectiveness in removing the unsintered powder, and shrinkage of the green part during post-processing affected the porosity of the actual part after sintering. The laser spot diameter of the DTM Sinterstation 2000 used in this study was ~450 µm.

The laser spot diameter of the DTM Sinterstation 2000 used in this study was ~450 µm. The resolution of laser spot would cause heating of the feedstock material on the part bed adjacent to the scanning area of the actual part, effectively reducing the designed pore size in the green part. Although the resolution could be improved with a smaller laser spot, the results mentioned in the current study are based on the laser spot available for the DTM 2000 Sinterstation. The apparent porosities of the sintered scaffolds were measured using Archimedes principle and are presented in Table 2, and the average pore sizes measured using optical microscope are presented in Table 3. The measured porosity results were all ~20% lower when compared to their original designs. Therefore, it has to be noted that along with the shrinkage factor, the difference in porosity factor should also be accounted for while fabricating porous parts using the indirect SLS fabrication process.

The measured porosity values were used to plot the graphs in the proceeding sections of the paper.Table 2. Differences in porosities: designed in CAD vs. measured after sintering.ArchitectureCubic Spherical X Diamond Gyroid50 33.38 31.53 27.79 33.85 34.31Porosity in %60 39.91 42.04 34.59 39.87 40.9170 53.88 49.02 43.17 46.73 49.480 65.76 60.58 54.98 61.37 59.46Table 3. Pore sizes of the sintered scaffolds at various porosities (in %).ArchitectureCubic Spherical X Diamond Gyroid50 500 500 400 800 80060 700 650 550 900 900Pore size (µm)70 950 800 700 1000 100080 1300 900 1200 1200 11003.2. Effect of ArchitectureFigure 2 shows the difference in the compressive strengths of the scaffolds made with different architectures at measured porosity levels of about ~30% and ~60%.

Effect of ArchitectureFigure 2 shows the difference in the compressive strengths of the scaffolds made with different architectures at measured porosity levels of about ~30% and ~60%. The vertical error bars represent the standard deviation of the measured compressive strength and the horizontal error bars represent the standard deviation in the measured porosity of the scaffolds. The scaffolds with a cubic architecture, and a porosity of ~33%, offered the highest compressive resistance among all of the scaffolds. This can be attributed to the struts in the z-direction which can carry a majority of the load, whereas, such a design feature is not present in any of the other architectures. The spherical and gyroid architectures offered a similar resistance which is followed by the diamond architecture.

The spherical and gyroid architectures offered a similar resistance which is followed by the diamond architecture. The scaffolds with the X architecture offered the leastresistance (7.4 MPa at ~28% porosity) compared to diamond architecture (7.2 MPa at ~34% porosity) during the compressive tests as the struts oriented at an angle of 45° to the base plane might not be ideal for compressive loads. Out of the two designs closely mimicking the trabecular bone, the gyroid seems to offer higher resistance than the diamond architecture because of the relatively thick (~1.3 mm for gyroid compared to ~0.9 mm for diamond) and wavy nature of the struts.

The compressive strengths of the scaffolds at the lower porosity level (~30%) fall in the range of the higher end of the range of compressive strengths for a human trabecular bone which ranges between 2 and 12 MPa [16], whereas the strengths measured for the scaffolds at higher porosity values (~60%) fall towards the lower range of the strength of human trabecular bone. The effect of different architectures at higher porosities (>60%) can be termed as not significant for load-bearing bone repair applications because of the little difference in compressive strengths of the scaffolds. Also, these higher porosity scaffolds during SBF tests have been shown to have a similar percentage reduction in strength, irrespective of architecture (discussed in detail in Section 3.4). This shows that the internal architecture design parameter could play a key role when fabricating implants at lower porosity levels for load-bearing applications.Figure 2.

This shows that the internal architecture design parameter could play a key role when fabricating implants at lower porosity levels for load-bearing applications.Figure 2. Compressive strengths for 13-93B3 scaffolds fabricated with five different pore geometries and at ~30% and ~60% porosity.3.3. Effect of PorosityFigure 3 shows the variation in compressive strengths of the five architectures at different porosities along with an optical image of the sintered scaffolds of diamond architecture at four porosity levels. Although the results show that the scaffold with cubic architecture provides the highest compressive strength at ~33% porosity, and the rate of reduction in strength is higher when compared to the rate of reduction in strength among all the scaffolds with different architectures. The scaffolds with the spherical architecture exhibited the smallest rate of reduction in compressive strength based on the slope of curve.

The scaffolds with the spherical architecture exhibited the smallest rate of reduction in compressive strength based on the slope of curve. The other three architectures namely, X, diamond, and gyroid have a similar rate of reduction in the compressive strengthswith respect to porosity. The reason for the higher reduction rate for the scaffolds with cubic architecture could be due to the design of the unit cube of this structure. To increase the porosity in the CAD model, the distance between the struts was increased, along with a reduction in the diameter of the strut, which may have resulted in a reduced resistance during compression tests.

In the case of scaffolds with the spherical architecture, the absence of struts to carry the load during compression tests results in the least effect on the rate of reduction with respect to increasing the porosity level.Finite element analysis (FEA) was performed to predict the compressive strengths of the scaffolds with the help of commercially available Abaqus software. For this purpose, the shrinkage and porosity reduction factors were incorporated in the CAD models. The files were modeled with 3D tetrahedral elements with sizes ~0.1 mm. The total number of elements for each of the models was kept in the range of 300,000 to 350,000. The ultimate strength of the 13- 93B3 material was determined experimentally by fabricating a solid dense cylinder and performing the compressive tests.

The ultimate strength of the 13- 93B3 material was determined experimentally by fabricating a solid dense cylinder and performing the compressive tests. The density of the solid dense cylinder was considered to have the material’s theoretical density (2.164 g/cc) because the same density will be achieved for all the SLS scaffolds as they go through the same heat treatment schedule. A Poisson’s ratio of 0.2 was assumed, which is typical for a bioceramic [17]. A 1.2% strain failure criterion was adopted to determine the failure of the specimen, which is typical for a porous bioactive glass specimen. The results in Figure 3 show that FEA can be used to achieve a fairly good estimate of the compressive strengths of the scaffolds fabricated with bioactive glass.Figure 3. Variation of compressive strength of scaffolds with different architectures at various porosities.3.4. SBF TestsThe scaffolds were mechanically tested in a wet state after immersion in SBF for one week.

SBF TestsThe scaffolds were mechanically tested in a wet state after immersion in SBF for one week. The compressive strength for each of the scaffolds was measured to be less than, or about, 2 MPa, which is towards the lower end of the range of compressive strength (2 – 12 MPa) of human trabecular bone [16]. The only exceptions were the scaffolds designed for 50% in CAD (actual measured porosities were ~33% and ~32% as shown in Table 2) with cubic and spherical architectures. The compressive strengths of these scaffolds after one week were measured to be ~4 MPa, which could be due to the relatively thick struts, smaller pores, and higher compressive strengths in their dry state. The reduction in the compressive strengths of the scaffolds with different architectures from ‘as-sintered dry state’ to ‘wet state’ after immersion in SBF for one week is shown in Figure 4.

The reduction in the compressive strengths of the scaffolds with different architectures from ‘as-sintered dry state’ to ‘wet state’ after immersion in SBF for one week is shown in Figure 4. Among the scaffolds which were designed for 80% porosity (actual measured porosities range between ~55% – 65%; see Table 2), irrespective of the architecture, the reduction in compressive strength in the wet scaffolds is more than ~90% when compared tothe as-sintered dry scaffolds. The ~90% reduction in strength could be due to the large pores measuring about ~1 mm in the scaffolds that were designed for 80% porosity (see Table 3). This allows the SBF to contact and react with the inside surface of the scaffold in an efficient manner. Among the scaffolds with lower porosities, the spherical architecture has the lowest reduction in compressive strength (57 – 65%), and the scaffolds with diamond architecture have the highest reduction in compressive strength (80 – 87%).

In comparison, the reduction in strength for the scaffolds fabricated using the Robocasting technique was reported to be around ~85% after 30 days [5]. This shows that the percentage reduction in the strengths of the SLS scaffolds is much higher when compared to the Robocasting scaffolds even when the scaffolds are made with the same bioactive glass. This could be because of the surface area that is offered for the scaffolds with complex architectures and also with the increased surface area associated with the SLS process itself (increased surface roughness of the struts) when compared to the Robocasting process (filaments/struts having smoother surface finish). This difference leads to the variable rates of degradation of the scaffolds when immersed in the SBF solution. The SBF results indicate that the SLS process could provide an opportunity, to a certain extent, to control the rate of degradation of scaffolds by selecting an appropriate architecture for the scaffold.

For example, the diamond architecture could be the choice for the scaffold fabricated with a silicate bioactive glass, which typically has a slower conversion rate, so as to augment the rate of degradation.Figure 4. The comparison of reduction in compressive strengths of scaffolds with different architectures and porosities after immersion in SBF for one week.After testing, the scaffolds were dried at room temperature, sputter coated with gold- palladium, and investigated using SEM. Specifically, SEM was used to look for any crystal like formations on the surface, which were typically formed within one week of the scaffolds’ immersion in SBF. Figure 5 shows SEM images of a representative “X” architecture scaffold taken out of the SBF after 7 days. The outer surface morphology of the SLS scaffold is shown in Figure 5(a). Figure 5(b) shows the fracture surface of the scaffold.

The outer surface morphology of the SLS scaffold is shown in Figure 5(a). Figure 5(b) shows the fracture surface of the scaffold. The reacted surface of the scaffold with SBF can be clearly distinguished with an unconverted 13-93B3 glass core as indicated in the image. A higher magnification image indicating the crystal like formations on the surface of the scaffold after reacting with the SBF is shown in Figure 5(c).Figure 5. SEM images of a scaffold after immersion in SBF for 7 days: (a) outer surface morphology, (b) fracture surface of the scaffold indicating the reacted glass (RG) on the outer surface and unconverted glass (G), (c) higher magnification image shows crystal-like formations on the surface of the scaffold.4. ConclusionsIt is concluded that among the borate-based bioactive glass scaffolds fabricated by the selective laser sintering (SLS) process, the cubic pore architecture provides the highest compressive strength at lower porosities (< 40%).

However, scaffolds with the cubic architecture also exhibit the highest rate of reduction in the compressive strength with increased porosity among all the architectures considered in this study. The effect of different architectures at higher porosities (> 60%) is not significant for load-bearing bone repair applications because of the small variability in the compressive strengths of the scaffolds. The compressive strengths at higher porosity levels (>60%), irrespective of the architecture, fall in the lower range of the compressive strength of a human trabecular bone. The finite element analysis has provided a close estimate of the compressive strengths of the scaffolds fabricated using the SLS process. The simulated body fluid tests indicate that the spherical architecture has the lowest percentage reduction and the diamond architecture has the highest percentage reduction in compressive strengths for scaffolds fabricated at similar porosities.

Therefore, the SLS process could provide an opportunity, to a certain extent, to control the rate of degradation of scaffolds by selecting an appropriate architecture for the scaffold.AcknowledgementsThe authors thankfully acknowledge the help of Priyank Madria for assistance and Mo-Sci Corp., Rolla, MO., for kindly supplying the 13-93B3 borate bioactive glass used in this work.References[1] Hench, L.L., Wilson, J., An Introduction to Bioceramics (Singapore: World ScientificPublishing Co. Pte. Ltd.) p 47 (1993).[2] Hench, L.L., The story of bioglass. J Mater Sci: Mater Med. 17, 967–978 (2006). [3] Rahaman, M.N., Day, D.E., Bal, B.S., Fu, Q., Jung, S.B., Bonewald, L.F., BioactiveGlass in Tissue Engineering. Acta Biomater. 7, 2355-2373 (2011).Fu, Q., Rahaman, M.N., Fu, H., Liu, X., Silicate, borosilicate, and borate bioactive glass scaffolds with controllable degradation rate for bone tissue engineering applications. I. Preparation and in vitro degradation. J. Biomed.

fect of Bimodal Powder Mixture on Powder Packing Density and Sintered Density in Binder Jetting of MetalsYun Bai, Grady Wagner, Christopher B. Williams Design, Research, and Education for Additive Manufacturing Systems Laboratory Department of Mechanical Engineering Virginia TechAbstractThe Binder Jetting Additive Manufacturing process provides an economical and scalable means of fabricating complex metal parts from a wide variety of materials. However, the performance metrics of the resulting sintered parts (e.g., thermal, electrical, and mechanical properties) are typically lower than traditionally manufactured counterparts due to challenges in achieving full theoretical density. This can be attributed to an imposed constraint on particle size and its deleterious effects on powder bed packing density and green part density.

This can be attributed to an imposed constraint on particle size and its deleterious effects on powder bed packing density and green part density. To address this issue, the authors explore the use of bimodal powder mixtures to improve the sintered density and material properties within the context of copper parts fabricated by Binder Jetting. The effectiveness of using bimodal powder mixtures in an effort to improve sintered density is studied in terms of particle size distribution and powder packing density.Key words: Binder jetting, Additive manufacturing, Copper, Sintering, 3D printing, Powder metallurgy1.Introduction1.1 Challenges in Binder Jetting of MetalsThe Binder Jetting Additive Manufacturing (AM) process can be used to fabricate metalparts by selectively ink-jetting a liquid binding agent into a powder bed, followed by post- process sintering of the printed green part.

In the green part creation stage, the binder droplets interact with the powder particles to form primitives that stitch together to form a cross-sectional layer. Once a layer is printed, a new layer of powder is spread by a counter-rotating roller on top of the previous layer which is then printed and stitched to the previous layer by the jetted binder. The layer-by-layer process is repeated to create the complete green part. The unbound loose powder in the bed that surrounds the part supports overhanging structures during the build, and can be removed after printing via compressed air. Once depowdered, the green part is placed in a high-temperature furnace to pyrolyze the binder and sinter the powder particles together through atomic diffusion in order to obtain final density and strength.Figure 1.

Green part printing process in Binder JettingAs Binder Jetting of metal functionally separates part creation from powder sintering, thecommon processing challenges found in direct-metal Additive Manufacturing processes – such as residual stresses imposed by rapid solidification of a melt pool, which lead to part warping and added anchors and/or heat sinks to geometry – are avoided (Mercelis and Kruth, 2006; Mumtaz et al., 2011; Shiomil et al., 2004). The ability to fabricate a part in a powder bed without the need for built anchors enables Binder Jetting to create large, geometrically complex parts without difficult post-process cleaning. Binder Jetting is also an inherently scalable technology as it does not require an enclosed chamber and expensive energy sources. In addition, as Binder Jetting does not use an energy beam to process material, it is well suited for optically reflective and thermally conductive metals, which can be challenging for Powder Bed Fusion processes.

For example, the feasibility of manufacturing high purity copper via Binder Jetting on ExOne 3D printers has been demonstrated in the authors’ previous work (Bai and Williams, 2015).The primary challenge in fabricating metal parts using Binder Jetting is in achieving afully dense product following the sintering post-process. Pores typically exist in sintered ceramics or metals fabricated in Binder Jetting (Chou et al., 2013; Zhang et al., 2014). For example, while the authors were able to create complex structures from gas atomized copper powder, the overall mechanical strength (116.7 MPa) and conductivity is constrained by the parts’ substantial porosity (85.5% dense parts). Part porosity is challenging to eliminate during sintering because of a low powder bed density and the inability to process ultra-fine powders. Loosely packed powders have less contacting points and large empty spaces between particles, which reduce available sintering and neck formation sites.

Loosely packed powders have less contacting points and large empty spaces between particles, which reduce available sintering and neck formation sites. The spreading of fine powders can be difficult in powder-based AM processes due to low flowability and powder agglomeration, but the presence of large particles in the powder significantly lowers the driving force for sintering (reduction of surface energy). As such, metal parts made by Binder Jetting are typically infiltrated with an infiltrant material in order to obtain full density.Pressure assisted sintering provides a viable solution to achieve pore-free density inBinder Jetting of metals. Green parts can be sintered to an initial density and then Hot Isostatic Pressing (HIP) can close the pore as a second treatment. The use of HIP has demonstrated to be able to approach full density in a slurry-based Binder Jetting of tungsten carbide (Kernan et al., 2007).

The use of HIP has demonstrated to be able to approach full density in a slurry-based Binder Jetting of tungsten carbide (Kernan et al., 2007). This strategy typically requires a high initial sintered density (usually over 93%) to enable capsule-free HIP so that complex shapes can remain intact. Thus, the goal of this work is to identify processing techniques with a focus on powder for improving sintered density of metal parts created by Binder Jetting to a value that is suitable for HIP.1.2 Improving sintered density via bimodal powder mixturesTo accomplish the goal of achieving a high initial sintered density, the authors look to themature Powder Metallurgy (PM) discipline as it is similar to Binder Jetting in that part sintering is separated from green part formation. One well-established theory in Powder Metallurgy to improve powder packing density is to use bimodal powder mixtures.

One well-established theory in Powder Metallurgy to improve powder packing density is to use bimodal powder mixtures. The increased powder packing density in bimodal powder mixtures, wherein the small particles fill the interstitial voids between coarse particles, has many benefits such as improved unfired property and less shrinkage given a certain sintered density. Zheng provides a review of the developed models anddemonstrated importance of the particle size distribution to obtain dense packing (Zheng et al., 1995).Compared with the certainty and reliable prediction in improving powder packing density, the sintered density improvement by bimodal powder is often complicated and can be unsuccessful. German developed a prediction of sintered density of bimodal mixtures and validated by a series of experiments with various materials (German, 1992).

German developed a prediction of sintered density of bimodal mixtures and validated by a series of experiments with various materials (German, 1992). Both models and practices showed the rapid sintering rate and enhanced sintering stress of small particles in sintering (Coble, 1973); however, added smaller particles typically bond to the large particles and offer little influence on the overall densification. The addition of large particles to fine powders may increase the packing density, but it will also hinder densification. In these two cases, a sintering stress develops in the powder mixtures and may inhibit densification, especially at lower sintering temperature and shorter time.

In these two cases, a sintering stress develops in the powder mixtures and may inhibit densification, especially at lower sintering temperature and shorter time. As a result, despite a successful increase in powder packing density, many sintering experiments involved with bimodal powder have failed to show a corresponding increase in the sintered density.Particle size distribution and bimodal mixtures have been explored in some powder-bed AM processes (SLS, SLM, etc.); however, the study of bimodal powder mixtures in the context of Binder Jetting is limited. In Powder Bed Fusion, particle size distribution plays an important role in increasing layer density (Karapatis and Egger, 1999); preventing balling phenomenon and achieving higher radiative heat flux in the powder bed (Zhou et al., 2009); and selecting the optimal printing parameters (Spierings et al., 2010; Spierings and G., 2009).

In Binder Jetting, Sachs studied the improved surface finish and printing primitives in unfired parts using bimodal powder distribution (Lanzetta and Sachs, 2003). Verlee explored the sintered density of stainless steel bimodal mixtures with various mixing ratios under one sintering condition. Like many Powder Metallurgy models have predicted, the experimental result did not show an improvement in sintered density by mixing large particles to the fine (Verlee et al., 2011).1.3 ContextWhile there is little demonstrated benefit of bimodal mixture in improving sintered density within Powder Metallurgy, the effect of bimodal powder mixture in Binder Jetting of copper is hypothesized to have a different impact on sintered density for the following reasons:1. The powder bed in Binder Jetting is much less dense than many traditional Powder Metallurgy especially for fine powders.

The powder bed in Binder Jetting is much less dense than many traditional Powder Metallurgy especially for fine powders. The low packing density is a main obstacle in achieving fully dense parts, which could have a strong impact on sintered density. 2. Binder Jetting of metals typically requires a large sintering temperature and duration,which could facilitate sintering stress relaxation, and thus denser final parts.3. Sintering behavior is material specific. Copper sinters well in PM contexts; however its study in the context of Binder Jetting is limited.Beyond improving sintering densification, the use of bimodal mixed powders in Binder Jetting is also expected to bring some additional benefits. As most metal parts in Binder Jetting undergo a large degree of shrinkage after sintering (without infiltration), there is a need to use high packing density powders to reduce shape distortion for a better dimensional control,especially for high precision parts.

In addition, a powder mixture containing coarse powders will have a decreased cost as compared to a powder bed composed of solely fine powders.This paper describes the authors’ efforts to validate the hypothesis that the use of abimodal powder mixture will improve final part density. This objective is achieved by processing copper parts made various particle size distributions and powder bed densities via bimodal powder mixtures. The authors’ experiments explore the effects of adding large/small particles into a base powder on the final sintered density under different sintering conditions. The experimental method is detailed in Section 2. A discussion of the results is presented in Section 3, and closure is offered in Section 4.

The experimental method is detailed in Section 2. A discussion of the results is presented in Section 3, and closure is offered in Section 4. While the results convey the effects of processing bimodal copper powder, the overall goal of this paper is to develop a fundamental understanding and methodology of material property improvement via powder optimization in Binder Jetting of metals.2.Experimental Method2.1 Powder selection and characterizationPowder characteristics such as morphology, size, and distribution affect final part quality.The authors have chosen to use gas atomized powders as they provide excellent packing density and are easily recoated.

As an additional benefit, spherical powder requires minimal binder to form necking between particles compared to irregularly shaped particles (Cima et al., 1992).Many models and experimental results have shown that high packing density can be achieved when the volume fraction of fine powder is between 0.2-0.4 (Zheng et al., 1995). In this paper, bimodal mixtures were created by mixing fine powder (15 m or 5 m) with coarse powder (75 m or 30 m) with a 73-27 weight ratios in a rotating drum. The powders used for mixing are listed in Table 1. Each mixture was mixed for 2 hours to ensure an even mixture. Laser scattering with a Horiba LA-950 was used to analyze particle size distribution (ASTM B822) of the powder mixtures.Table 1.

Each mixture was mixed for 2 hours to ensure an even mixture. Laser scattering with a Horiba LA-950 was used to analyze particle size distribution (ASTM B822) of the powder mixtures.Table 1. Powder descriptionPowder name D10 D50 D90 𝐻2 Loss% 75 μm powder 58.0 77.0 101.5 30 μm powder 15.0 30.0 37.5 28 15 μm powder 8.0 17.0 9.0 5 μm powder 0.65 5.5N/A 0.40 0.30 0.652.2 Powder bed analysisAs bimodal mixtures are hypothesized to significantly affect green part density, and in Binder Jetting the green density largely depends on the powder bed density, the authors employed a variety of methods for evaluating powder bed density. Apparent density, which should be the lower density threshold in Binder Jetting, was measured using a Hall flow meter (ASTM Standard 212). Tap density, which should be the upper threshold, was measured using a tapping apparatus(ASTM Standard 527).

Tap density, which should be the upper threshold, was measured using a tapping apparatus(ASTM Standard 527). The actual powder bed density should be in between the apparent and tap density because of the compaction of the spreader; however, this density is usually difficult to be directly measured. Inspired by some other powder-bed AM processes, in this work a cup was first printed into the powder; powder bed density was then calculated by measuring the mass of powder contained in the cup dividing the cup volume.Apparent and tap density is convenient to measure and usually comes with high accuracy; therefore these two densities can be used for powder screening in Binder Jetting, especially when packing density is the main concern. In addition, the ratio of apparent and tap density (Hausner ratio) can assess powder flowability – the lower this ratio, the better the flowability.

In addition, the ratio of apparent and tap density (Hausner ratio) can assess powder flowability – the lower this ratio, the better the flowability. Powder packing characteristics can be assessed by directly measuring the powder bed density with a printed cup or measuring green part density. The printed cup method is preferred as it eliminates the binder effect; however, the measurement can lose its accuracy when ultra-fine powders (non- free flowing) are used as they prove to be difficult to completely remove from the cup.2.3 Printing process parametersAn off-the-shelf standard metal binder (PM-B-SR2-05) from ExOne was used for allexperiments as it leaves minimal binder residue after sintering and is compatible with both the Binder Jetting machine (ExOne R2) and copper powder used in this study.

18mm x 6mm x 3 mm test coupons were printed for characterization of the various powder mixtures.To ensure successful spreading, the layer thickness must be larger than the largestparticle, and it is recommended that the layer thickness should be at least three times the layer thickness to acquire a higher packing density and smoother surface finish (Utela et al., 2008). In addition, the layer thickness cannot be larger than the radius of a primitive to ensure that subsequent layers are bound together. In this paper all samples are printed with an 80 µm layer thickness.Saturation ratio is the ratio of the amount of void space in the powder bed filled withbinder to the total amount of void space. Too much binder will permeate through the powder and bind extra powder, which leads to a green part growth and poor surface finish.

Too much binder will permeate through the powder and bind extra powder, which leads to a green part growth and poor surface finish. However, if saturation ratio is set too low, printed binder will not penetrate deep enough to bind layers, which can cause porosity, anisotropic shrinkage, and part delamination. An accurate binder drop volume measurement was made before each print by jetting a known amount binder droplets onto a testing substrate and measuring the total mass. As the fractional volume of void space is different for each explored powder combination, a unique packing density is entered into the process control software for each bimodal powder, and is based on the powder bed density measurements outlined in Section 2.1. The amount of jetted binder is adjusted accordingly to obtain the predetermined saturation value.

The amount of jetted binder is adjusted accordingly to obtain the predetermined saturation value. Based on the considerations stated above and an observation of the primitive size in each powder, a saturation of 150% is used for the 5 μm powder and 100% for the rest. The 5 μm powder has displayed a small primitive size and large total surface area thus requires extra binder for good bonding.2.4 Post-processing and sinteringThe chosen PM-B-SR2-05 binder typically requires curing at 190 °C for 2 hours after printing to increase green part strength. All sintering cycles used in this work feature constant heating/cooling rate (3 °C/min) and contain an isotherm at 450 °C for 30 min to facilitate binder burn-out.

All sintering cycles used in this work feature constant heating/cooling rate (3 °C/min) and contain an isotherm at 450 °C for 30 min to facilitate binder burn-out. To explore the hypothesized different response of bimodal powder mixtures under various sintering conditions, a full factorial design containing two levels of sintering temperature (1020 °C and 1060 °C) and two levels of sintering time (30 min and 120 min) was used for sintering the printed testing coupons.To measure sintered density, the immersion method with Archimedes principle was useddue to the porosity of the part (ASTM Standard 962). This method is used for samples with surface connected porosity by filling the pores with a known content of oil.

This method is used for samples with surface connected porosity by filling the pores with a known content of oil. Shrinkage is calculated by comparing the green dimensions to the sintered dimensions.3.Results and Discussion3.1 Powder bed density, green part density, and flowability of bimodal powder mixturesTables 2 summarizes the powder and powder mixtures with a decreasing order of median size, and has shown the apparent/tap density improvement in bimodal mixtures. For example, by introducing a small portion of large particles (30 μm or 15 μm) to the 5 μm powder (powder 9), the apparent densities of the resultant powder beds (powder 7 and 8) have improved by 12.7% and 5.6% and the tap density has improved by 5.6% and 4.9%. The increase in apparent density is higher than that in tap density, which is particularly useful for Binder Jetting as its powder compaction effect is limited.

The increase in apparent density is higher than that in tap density, which is particularly useful for Binder Jetting as its powder compaction effect is limited. Except for the large particle size ratio powder mixtures (75+15 μm mixture) where a dual peak can be observed in the particle size distribution curve, the powder density increase in most mixtures are achieved by a shifted median size and widened size distribution. The relatively small large-to-small particle size ratio in this work (lower than 6 ) limited the increase in powder packing density; however it is necessary to keep this ratio small so that no extra-large particles are introduced to the mixtures that will be a detriment to sintering.The Hausner ratio has shown the improved powder flowability (smaller ratio) in allbimodal mixtures (Table 2). The improved flowability is critical in getting smooth and dense powder layers in the recoating process of Binder Jetting.Table 2.

In Figure 2, the value of powder bed density (obtained by measuring powder in a printed cup; Section 2.2) is measured as in between of apparent density and tap density in each powder. The powder bed density is more similar to the apparent density for fine powders, while it is more similar to the tap density for high flowability powders (75 µm and 30 µm). Figure 2 also compares the green density with the apparent and tap densities of selected powder mixtures. The green density of the printed parts is slightly lower than the apparent density for most powders (except for powder 9). This is believed to be caused by the binder spread effect at the edge of the printed parts that glues extra powder particles, which result in a green part dimensional expansion over the original design.All density measurement methods have shown noticeable increases in powder density inthe printed parts.

In Figure 2, it is seen that a better powder bed density is achieved by either mixing small particles (5 µm) or mixing large particles (75 µm) into the 15 µm powder, with an increase of 5.7% and 16.2% respectively. Figure 2 also shows that the mixtures containing 5 µm powders improve the green density by 3.0%-9.4% when compared to the pure 5 µm powder. This result shows a means for increasing powder bed density while still retaining most of the fine powders, which are important for sintering performance.y t i s n e d r e d w o P80% 70% 60% 50% 40% 30% 20% 10% 0%57.20%52.20%48.20%53.40%51.40%47.00%44.00%27.017.517.010.88.35.85.575 μm (w.t.73%)+15 μm30 μm (w.t.73%)+5 μm15 μm15 μm (w.t.73%)+5 μmMedian size (µm)5 μm (w.t.73%)+30 μm5 μm (w.t.73%)+15 μm5 μmMeasured powder bed densityApparent densityTap densityGreen part denstiyFigure 2.

Comparison of apparent density, tap density, green density (with density numbers shown in the graph), and measured powder bed density (only for the coarse powders to ensure measurement accuracy) in Binder Jetting3.2 Packing density and particle size effect on sinteringTo explore the effects of bimodal mixtures on part densification during sintering, a non-mixed powder (15 μm) and three powder mixtures (15 μm+5 μm powder, 75 μm+15 μm powder, 30 μm+5 μm powder) were printed into testing coupons and sintered. The mixing ratio was held constant for this section (73% coarse powder). As noted in Section 3.1, compared with the 15 μm powder, the 15 μm+5 μm powder introduces fine particles into the mixture and has increased powder bed density by 5.7%; the 75 μm+15 μm powder brings large particles into the mixture and has improved the powder bed density by 16.2 %.Samples were printed using the same printing parameters as described in Section 2.3.

Theisotherm temperature and time in the sintering heating schedule was varied at two levels (1020 °C and 1060 °C, and 30 minutes and 120 minutes), as indicated in Figure 3. In each quadrant, the sintered density (lower section) and densification (upper section) of each sintering cycle are compared across the powder mixtures. The densification is represented by the density gain from powder bed density to sintered density.1020 °C1060 °C30 min(a)(b)120 min(c)(d)Figure 3. Sintered density and densification comparison across the powder under various sintering conditionsThe parts sintered under the least sufficient sintering condition (1020 °C for 30 minutes)achieved less densification than higher isotherm temperatures and longer holds.

In addition, Figure 3 (a) shows that when large particles are present in Binder Jetting (75+15 µm and 30+5 µm powder mixtures), the sintered density can be lower than the powder bed density (the parts still densified as the actual green density is lower than the measured powder bed density). Under this condition, powder bed density still plays an important role in determining sintered density as the sintering densification is limited.An increase in powder bed density by introducing extra-large particles has failed to effectively improve the sintered density. This is demonstrated by comparing the 75+15 µm powder mixture with the 15 µm powder in Figure 3(b), (c) and (d). Figure 4(b) shows the necking has a tendency to form between 15 µm and 75 µm particle. There is limited sign of direct necking and merging of 75 µm particles due to the large particle radii.

Figure 4(b) shows the necking has a tendency to form between 15 µm and 75 µm particle. There is limited sign of direct necking and merging of 75 µm particles due to the large particle radii. The sintered part is composed of a rigid skeleton of 75 µm particles, wherein the sintering of fine powders have littlecontribution to the overall densification. This phenomenon is not alleviated by the sintering temperature studied in this work. This result is similar to what has been found in traditional Powder Metallurgy (German, 1992).From these results, it is clear that an increase in powder bed density without shiftingmedian particle size is an effective method to improve sintered density. This is demonstrated by comparing the 30+5 µm powder mixture with the 15 µm powder. Compared with the 15 µm powder (median size of 17.0 µm), the 30+5 µm powder mixture has a very similar median particle size (17.4 µm) but a wider distribution, which results 9.2% higher powder bed density.

This difference results in a 7.5%-11.7% sintered density depending on the sintering condition (Figure 3).The highest sintered density (87.1%) occurred in 30+5 µm powder samples when sinteredat 1060 °C for 120 minutes. The powder mixture provides more contact point as available neck formation sites, thus small particles are able to “glue” neighboring large particles and form an integrated structure by surface reduction (Figure 5). Despite the significant increase in sintered density, the volumetric shrinkage in 30+5 µm powder only increased by 2.4% than in the 15 µm powder (Figure 6), which is another benefit of powder mixing. The linear shrinkage is anisotropic but displays a similar trend.(a) (b) Figure 4. Sintered structures of the primitives in (a) 15+5 µm powder and (b) 75+5 µm powder, sintered at 1060 °C for 120 minutes(a) (b)Figure 5.

Sintered structures of the primitives in (a) 15+5 µm powder and (b) 75+5 µm powder, sintered at 1060 °C for 120 minutes(a) (b)Figure 5. Sintered structures of the primitives in 30+5 µm powder, sintered at 1060 °C for 120 MinutesSimilarly, the mixing of 5 µm powder into 15 µm powder is capable of increasing sintered density as both powder bed density and particle size is improved. 15+5 µm powder has the highest density gain among the powders tested as it has the smallest median size. Like 30+5 µm powder, interconnected sintered agglomerates are observed in Figure 4(a).(a) (b) Figure 6.

Like 30+5 µm powder, interconnected sintered agglomerates are observed in Figure 4(a).(a) (b) Figure 6. (a) Volumetric shrinkage and (b) X, Y and Z linear shrinkage of the powders sintered at 1060 °C for 120 minutes4.ConclusionThe bimodal powder mixtures in the context of Binder Jetting of copper is capable of improving spreading, powder bed density, and sintered density compared with the constituent powders.The strategy of increasing powder bed density and flowability via bimodal powdermixtures is particularly effective for Binder Jetting. By adopting bimodal powder mixtures, the powder flowability is improved by up to 15% (30+5 µm powder; Section 3.1). Despite the relatively small particle size ratio (3-6 when 15 µm powder or 30 µm powder is mixed with 5 µm powder) used in this work, the powder bed density is able to be improved by up to 16.2% (Section 3.1).

However this dramatic increase in the powder bed density has failed to improve the sintered density when large particles (75 µm powder) are present as they formed a rigid skeleton which limited the contribution of fine powders to overall densification. Widening particle size distribution while not shifting median particle size via bimodal mixtures (e.g., 30+5 µm powder) has demonstrated to be a viable strategy and has shown an increase of sintered density up to 12.3 % depending on the sintering condition (Section 3.2).This work provides a framework for powder optimization in terms of achieving high sintered density in Binder Jetting of bimodal metal powders. The authors look to explore the effects of different printing parameters and optimize sintering profiles on final part density in future work.5.AcknowledgementThis material is based upon work supported by the National Science Foundation underGrant No. #1254287.

#1254287. Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the views of the National Science Foundation. The authors acknowledge technical support provided by the ExOne Co. The authors also would like to thank NCFL of Virginia Tech, Dr. Carlos Suchicital and Jeremy Beach for their assistance. Special thanks to Dan Brunermer (ExOne Co.) and Joe Strauss (HJE Co.) for the insight on Binder Jetting and Powder Metallurgy.6.ReferenceBai, Y. and Williams, C.B. (2015), “An exploration of binder jetting of copper”, RapidPrototyping Journal, Vol. 21 No. 2, pp. 177–185.Chou, D.T., Wells, D., Hong, D., Lee, B., Kuhn, H. and Kumta, P.N. (2013), “Novel processing of iron-manganese alloy-based biomaterials by inkjet 3-D printing”, Acta Biomaterialia, Acta Materialia Inc., Vol. 9 No. 10, pp. 8593–8603.Cima, M.J., Lauder, A., Khanuja, S. and Sachs, E.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEFFECT OF BUILD ORIENTATION ON FATIGUE PERFORMANCE OF TI-6AL-4V PARTS FABRICATED VIA LASER-BASED POWDER BED FUSIONBrian Torries, Nima Shamsaei*, Scott ThompsonLaboratory for Fatigue and Additive Manufacturing Excellence (FAME), Department ofMechanical Engineering, Auburn, AL 36849Corresponding author: shamsaei@auburn.eduAbstractThe effects of build orientation on the fatigue behavior of additively-manufactured Ti-6Al- 4V using a Laser-Based Power Bed Fusion (L-PBF) process is investigated. Ti-6Al-4V rods were manufactured in vertical, horizontal, and 45º angle orientations. The specimens were then machined and polished along the gage section in order to reduce the effects of surface roughness on fatigue behavior.

The specimens were then machined and polished along the gage section in order to reduce the effects of surface roughness on fatigue behavior. Fully-reversed strain-controlled uniaxial fatigue tests were performed at various strain amplitudes with frequencies adjusted to maintain an average constant strain rate throughout testing. Results indicate slight variation in fatigue behavior of specimens fabricated in the different orientations investigated. Fractography was conducted using scanning electron microscopy after mechanical testing in order to investigate the crack initiation sites and determine the defect responsible for the failure.

Fractography was conducted using scanning electron microscopy after mechanical testing in order to investigate the crack initiation sites and determine the defect responsible for the failure. The experimental program utilized and results obtained will be presented and discussed.Keywords: Fatigue, Fractography, Build Orientation, Additive Manufacturing, Ti-6Al-4VIntroductionAdditive manufacturing (AM) is a promising avenue of part fabrication due to the ability to fabricate parts in a layer by layer fashion and to create geometries difficult or impossible to achieve by traditional subtractive methods. There are two broad categories of laser-based additive manufacturing: Laser-powder bed fusion (L-PBF) and direct laser deposition (DLD) [1]. L-PBF processes, such as selective laser melting (SLM), involve the deposition of thin layers of metallic powder over a build substrate, with a laser moving in a predetermined pattern in order to create a layer.

After this occurs, the substrate is incremented down, and the process repeats until the part is fully formed.Layer orientation has been shown to significantly affect the fatigue properties of additively- manufactured materials [2]. Fatigue testing of L-PBF stainless steel 316L by Shrestha et al. [3] showed that specimens fabricated in a diagonal orientation displayed lower fatigue lives than those fabricated in the vertical direction, which in turn exhibited lower fatigue lives than horizontally fabricated specimens. Previous study by Yadollahi et al. [4] has shown that the layer orientation115of 17-4 PH steel parts produced using L-PBF can have an effect on the tensile and mechanical properties of the part. It was found that specimens printed horizontally had better fatigue properties and more elongation to failure than vertically built specimens did.

It was found that specimens printed horizontally had better fatigue properties and more elongation to failure than vertically built specimens did. The relative orientation of defects with respect to the loading direction lead to variation in the size of the projection of the void onto the plane perpendicular to the applied load. As the size of these projections increased, the resultant stress concentration also increased, resulting in lower fatigue strength of the specimens [4].Results from Edwards and Ramulu have shown that this is true in Ti-6Al-4V as well, with specimens machined horizontally from L-PBF fabricated Ti-6Al-4V displaying longer fatigue lives than the vertically machined specimens [5]. In another study of Ti-6Al-4V fabricated via SLM [6], specimens printed vertically had better fatigue performance than cast and annealed Ti- 6Al-4V. The SLM Ti-6Al-4V also possessed higher yield stress and ultimate tensile strength as compared to published values in the ASM Handbook.

The SLM Ti-6Al-4V also possessed higher yield stress and ultimate tensile strength as compared to published values in the ASM Handbook. This improved fatigue performance was attributed to the presence of fine lamellar structures and acicular microstructure in the SLM Ti- 6Al-4V [6]. Other studies attributed the difference in fatigue properties between different build orientations to orientation of prior-β boundaries [7]. Porosity is also a major contributing factor to the fatigue strength of L-PBF [8-10], with porosity drastically reducing high cycle fatigue life [8] and porosity caused by insufficient energy input (i.e. lack of fusion) during SLM fabrication having a much greater impact on mechanical behavior than porosity caused by super-optimum energy input [9].

lack of fusion) during SLM fabrication having a much greater impact on mechanical behavior than porosity caused by super-optimum energy input [9]. Therefore, the relative orientation of pores with respect to the loading direction may cause anisotropy in mechanical and fatigue performance of additively-manufactured parts.In this study, the effects of the build orientation on the fatigue properties of L-PBF Ti-6Al- 4V is investigated. Using an EOS M290, rod stock was fabricated in three batches of different orientations: horizontal, vertical, and 45° with respect to the build plate. This rod stock was then machined into fatigue specimens and subjected to fatigue testing, with fractographic analysis following failure. Finally, conclusions were drawn based upon these results.Experimental SetupFor the testing conducted in this study, 30 total specimens were fabricated.

Finally, conclusions were drawn based upon these results.Experimental SetupFor the testing conducted in this study, 30 total specimens were fabricated. Three batches of ten L-PBF rods were fabricated with a batch each in the vertical, horizontal, and 45° orientation (i.e. diagonal). All specimens were fabricated using an EOS M290 machine, gas atomized Ti-6Al- 4V powder supplied by LPW Technology Inc. (particle size 15 µm - 45 µm), and the default performance Ti-6Al-4V process parameters provided by the equipment manufacturer. Default support structures of 3 mm were used for each specimen in order to facilitate easy removal from the building substrate.After fabrication, the specimens were subjected to an annealing process consisting of a one-hour soak at 700°C in an argon atmosphere, followed by cooling to room temperature via free convection with the hope of removing any residual stresses from the fabrication process.

After annealing, the specimens were removed from the substrate and machined into round fatigue specimens with straight gage sections of 15 mm length and 3.75 mm diameter, compliant with ASTM standard E606 [11]. Once the machining was completed, the surfaces were polished with progressively finer grits of sandpaper in order to reduce the surface roughness in the gage section. The specimens were then subjected to fully-reversed, strain-controlled uniaxial fatigue testing, ranging from 0.004 to 0.01 mm/mm, conducted using an MTS Landmark servohydraulic test system with 100 kN load cells. Test frequencies were adjusted to maintain an average constant strain rate among all tests. An MTS Extensometer was used in order to measure and control the116strain applied to the specimens. Acrylic was applied to the gage section in order to prevent slippage of the extensometer and prevent scratching of the specimen surface.

Acrylic was applied to the gage section in order to prevent slippage of the extensometer and prevent scratching of the specimen surface. Fracture surfaces of failed specimens were investigated via field emission gun scanning electron microscope (FEG-SEM) in order to determine the crack initiation sites and crack growth behavior, as well as to document any porosity or lack of fusion areas on these fracture surfaces.Experimental Results and DiscussionFatigue Behavior Strain-life curves derived from the fatigue data collected during the strain-controlled testing of the L-PBF Ti-6Al-4V specimens are presented in Figure 1. As can be seen from the figure, the specimens possessed similar average fatigue lives regardless of build orientation. Therefore, the build orientation appears to have not had a significant impact on the fatigue responses of the specimens fabricated using EOS M290.

Therefore, the build orientation appears to have not had a significant impact on the fatigue responses of the specimens fabricated using EOS M290. While this is in disagreement with the study by Edwards and Ramulu, in which it was found that build orientation did have a noticeable effect on the fatigue performance of SLM Ti-6Al-4V [5], the samples used in their study contained much more porosity than those fabricated and tested in this study. This difference in porosity could account for the difference in findings, as the porosity contained within SLM specimens has been shown to affect fatigue life [4, 8-10].Figure 1. Comparison of strain-life fatigue behavior for diagonal, horizontal, and vertical L-PBF Ti-6Al-4V specimens.The lower levels of porosity contained within the specimens in this study led to less pore orientation effect, resulting in a lessening of the impact build orientation can possibly have on fatigue life.

While there were no runout tests, specimens in lower strain amplitudes did exhibit fatigue lives of greater than 2,000,000 reversals. There is generally less scatter in the low cycle regime as opposed to the high cycle regime for all orientations; this is due to the greater effect of porosity in the high cycle regime [12]. The diagonal specimens had the least amount of scatter of the tested specimens.117Fractography After the fatigue testing was completed, fractography was conducted on the fracture surfaces. Very little porosity was found during the fractographic investigation, with the porosity that was seen on the fracture surfaces serving as the crack initiation site. Inspection of the fracture surfaces showed that the cracks generally initiated at either pores or lack of fusion areas. Pores had smooth surfaces, usually with either spherical or ovoid geometries.

Inspection of the fracture surfaces showed that the cracks generally initiated at either pores or lack of fusion areas. Pores had smooth surfaces, usually with either spherical or ovoid geometries. Lack of fusion areas ranged in size from small to large enough to be visible without the aid of a microscope, and had a smooth surface on one side of the fracture surface and a rough, powder covered surface on the sibling fracture surface. These smooth and rough surfaces of the lack of fusion area were on the top and bottom, respectively, during the fabrication process.Figure 2 shows a representative fracture surface and crack initiation site for the vertical specimens. These surfaces exhibited crack initiation at a subsurface pore and evidence of crack propagation, with clear final fracture and ductile shear in the overload area.

These surfaces exhibited crack initiation at a subsurface pore and evidence of crack propagation, with clear final fracture and ductile shear in the overload area. While the horizontal specimen fracture surfaces displayed similar properties, as seen in Figure 3, the fracture surface was smoother than that of the vertical specimens. This could be due to the pore that served as the crack initiation site in the horizontal specimen; it was larger than the pore seen in Figure 2 (83 µm versus 12 µm), it was also more uniformly shaped and much further away from the surface of the specimen. This could have lessened the severity of the pore’s influence, allowing for longer crack propagation in horizontal specimens. Fracture surfaces of the diagonal specimens, shown in Figure 4, exhibited much less crack propagation and rougher fracture surfaces as compared to the vertical and horizontal specimens.Figure 2.

Fracture surfaces of the diagonal specimens, shown in Figure 4, exhibited much less crack propagation and rougher fracture surfaces as compared to the vertical and horizontal specimens.Figure 2. Overview and crack initiation site of a vertical L-PBF Ti-6Al-4V specimen.118Figure 3. Overview and crack initiation site of a horizontal L-PBF Ti-6Al-4V specimen.Figure 4. Overview and crack initiation site of a diagonal L-PBF Ti-6Al-4V specimen.ConclusionsThe goal of this study was to quantify the effects of build orientation on the fatigue performance of L-PBF produced Ti-6Al-4V. In order to accomplish this, fully-reversed, strain- controlled fatigue testing was conducted on specimens machined from rod stock printed in the vertical, horizontal, and 45° angle (i.e. diagonal) orientations using an EOS M290 machine. After testing, fractography was conducted on the resultant fracture surfaces. The following conclusions can be made based upon the results:1.

diagonal) orientations using an EOS M290 machine. After testing, fractography was conducted on the resultant fracture surfaces. The following conclusions can be made based upon the results:1. Layer orientation did not have a significant effect on the fatigue response of the L-PBF Ti- 6Al-4V, with all specimens displaying nearly identical fatigue properties. This could be due to the low levels of porosity observed in these specimens.1192. There were no runout tests. While specimens tested at lower strain amplitudes did exhibit fatigue lives greater than 2,000,000 reversals, the tests did fail eventually.3. Cracks were observed to initiate at pores or lack of fusion areas just below the surface in all specimens.4.

Cracks were observed to initiate at pores or lack of fusion areas just below the surface in all specimens.4. While there was little difference in the fatigue behavior of the specimens and crack initiation sites, the fracture surfaces were radically different, with varying levels of roughness and crack propagation.AcknowledgmentThis material is based upon work supported by the National Science Foundation (NSF)under Grant No. 1657195.References[1] Thompson, S. M., Bian, L., Shamsaei, N., Yadollahi, A., “An Overview of Direct Laser Deposition for Additive Manufacturing; Part I: Transport Phenomena, Modeling and Diagnostics,” Additive Manufacturing, Vol 8, 2015, pp. 36-62. [2] Shamsaei, N., Yadollahi, A., Bian, L., Thompson, S. M., “An Overview of Direct Laser Deposition for Additive Manufacturing; Part II: Mechanical Behavior, Process Parameter Optimization and Control,” Additive Manufacturing, Vol. 8, 2015, pp. 12-35.

viewed PaperEffect of Build Orientation on Residual Stress and Microstructure in Inconel 625 Fabricated via Laser Powder Bed FusionM. Andurkar1, B. C. Prorok2, S. M. Thompson*11 Alan Levin Department of Mechanical and Nuclear Engineering, Kansas State University, Manhattan, KS 66502, United States of America2 Department of Materials Engineering, Auburn University, Auburn, AL 36849, United States of AmericaCorresponding author: smthompson@ksu.eduAbstractThe reliability of parts produced by Laser Powder Bed Fusion (L-PBF) is still not at a great acceptance level. One of the major defects inherited in parts fabricated from L-PBF is a high level of residual stress. In this study, two build orientations i.e., vertical and diagonal, were used to fabricate Inconel 625 specimens to observe its effects on the residual stress magnitude and grain growth. A novel, Cos-α X-ray diffraction method was used to measure residual stress values along the top surface of the samples.

A novel, Cos-α X-ray diffraction method was used to measure residual stress values along the top surface of the samples. Electron Backscattered Diffraction (EBSD) and kernel average misorientation (KAM) maps were employed to explain residual stress trends and differences between samples. Results indicate that the as-printed vertical sample possessed a higher tensile residual stress (77 ± 15 MPa) compared to the diagonally-printed sample (52 ± 12 MPa). The KAM map of the as-printed vertically oriented sample showed more pronounced local misorientations caused by dislocations compared to the diagonally-printed sample.Keywords: Selective Laser Melting, Nickel Superalloy, EBSD, Misorientation, Microstrain1. IntroductionThe Laser powder bed fusion (L-PBF) process is one of the seven different categories of additive manufacturing (AM) technologies according to the ASTM 52900:2021 [1].

IntroductionThe Laser powder bed fusion (L-PBF) process is one of the seven different categories of additive manufacturing (AM) technologies according to the ASTM 52900:2021 [1]. The importance and interest of this technology has proliferated in various industrial sectors, in particular aerospace, automotive, nuclear, biomedical. L-PBF offers several advantages like fabricating complex geometries, rapid design to manufacture cycles, reduction of wasted materials, and an ability to tailor properties of end-user components during the fabrication process. The L- PBF process consists of melting consecutive powder layers by a focused laser beam guided by 3D- computer aided design (CAD). The presence of support structures and the powder bed itself allows L-PBF to be capable of building complex shapes of metal parts that would be challenging to fabricate using conventional methods.

The presence of support structures and the powder bed itself allows L-PBF to be capable of building complex shapes of metal parts that would be challenging to fabricate using conventional methods. Laser employed manufacturing methods involve large thermal gradients in the vicinity of the targeted region due to high heat flux. The melted layers are subsequently subjected to extremely high cooling rates (around 105–106 oC/s) that generate large non-uniform thermal gradients. This large thermal gradients precede the formation of residual stress also known as ‘hidden stress’.

This large thermal gradients precede the formation of residual stress also known as ‘hidden stress’. Residual stress left unmanaged in the finished product can lead to part distortion, delamination, parts built outside geometric tolerances, increased risk of crack formation, and reduced fatigue performance [2].1800Nickel based superalloys are known to be notoriously challenging to use in machining, as wells as conventional casting or forging methods due to its high hardness, low thermal conductivity, and high work hardening rate [3], [4]. Inconel 625 (IN625) is a nickel based solid solution strengthening superalloy which derives its strengthening ability largely from the presence of molybdenum (Mo) and Niobium (Nb) elements. IN625 finds its application in different industry sectors such as aeronautics, aerospace, nuclear, chemical and petrochemical due to superior properties like high-temperature strength, high creep resistance, and corrosion resistance [5], [6][7].

Due to challenges posed by using subtractive manufacturing methods and the recent bombarding developments of AM technology, fabrication of IN625 components by AM according to industrial standards has emerged as a promising approach. However, IN625 parts produced by methods such as L-PBF still cause challenges like the formation of residual stress. A careful selection of process parameters, scan strategy, build orientation is needed to fabricate parts with optimized microstructure, density, and mechanical properties.In the past, studies have been conducted to understand the effects of different scan strategies and build orientations on the formation of residual stress in different materials. Pant et al. mapped residual stress in Inconel 718 produced by L-PBF in three orientations (vertical, diagonal, horizontal). They observed the vertical sample inherited the highest tensile residual stress followed by the diagonal and horizontal samples [8].

They observed the vertical sample inherited the highest tensile residual stress followed by the diagonal and horizontal samples [8]. Some studies reported the greatest residual stress was generated perpendicular to scan direction [9], while others have observed the greatest residual stress parallel to scan vectors [10]. In addition, unidirectional scan vectors may give rise to higher anisotropic residual stress compared to checkerboard or alternating scanning strategies since longer scanning paths introduce heterogenous thermal gradients [11]. To understand the effects of different scan strategies and build orientation on residual stress formation in AM parts, accurate and reliable residual stress measurement techniques are critical.Measurement of residual stress can be carried out using destructive or non-destructive methods. Non-destructive testing (NDT) category involves X-ray, neutrons, synchrotron radiation, and magnetic strain[12].

Non-destructive testing (NDT) category involves X-ray, neutrons, synchrotron radiation, and magnetic strain[12]. X-ray diffraction (XRD) is one of the most matured technologies among all NDT methods. XRD is used to measure residual stress on external surface and sub-surface with high accuracy. XRD typically relies on using the sin²Ψ method for calculating stresses. The XRD- sin2Ψ stress method begins with XRD scans along series of defined incident angles (Ψ) over a pre- determined diffraction angle. From the scans performed at each incident angle, the selected diffraction peak position must be accurately verified, i.e., the 2θ value must be measured from a very wide-ranging and sometimes unevenly shaped peak. The obtained 2θ values are then used to execute the linear regression to obtain the slope and intercept values to be used for the stress calculation.

The obtained 2θ values are then used to execute the linear regression to obtain the slope and intercept values to be used for the stress calculation. Line detectors are usually used in the sin2 Ψ method which captures a small part of the Debye-Scherrer (D-S) ring at every incident angle (Ψ). The D-S ring is created by a diffraction pattern occurring due to Braggs law when X-rays are incident on a polycrystalline material. To eliminate the extensive process time to obtain data from this method, a novel method of Cos α was developed in 1978. The Cos α method employs the entire D-S ring produced from single exposure of X-rays at a single incident X-ray beam angle captured by a 2-D detector. The Cos α method has several advantages over the traditional sin²Ψ method such as simple optical setup and faster in- plane normal and shear stress calculations [13].

The Cos α method has several advantages over the traditional sin²Ψ method such as simple optical setup and faster in- plane normal and shear stress calculations [13]. The accuracy of this novel method was tested against the traditional sin²Ψ method and it was found to provide reliable results.In this current work, the effect of build orientation i.e., vertical and diagonal on residual stress value and grain morphology in Inconel 625 produced by laser powder bed fusion additive1801manufacturing process. A novel method was used to measure residual stress which is proven to be easier, quicker, and reliable. A comparison of residual stress between two build orientations was discussed along with grain morphology and kernel average misorientations. This study will correlate the residual stress results with grain texture and KAM maps.2. Methods2.1.

This study will correlate the residual stress results with grain texture and KAM maps.2. Methods2.1. AM processGas atomized virgin Inconel 625 powder was procured from Carpenter Technologies with mean particle diameter size of 31.3 µm. A Concept Laser Mlab 100R commercial L-PBF system was used to fabricate Inconel 625 samples. The L-PBF system works with 100 W fiber laser at a wavelength of 70 nm. The samples were fabricated in an enclosed argon atmosphere to avoid any oxidation of alloy elements and provide unhampering flow of powder during the process. The elemental composition of virgin Inconel 625 powder used during L-PBF process was following the ASTM F3056-14 (21) standard. The elemental composition (in % wt) in the L-PBF powder and standard is presented in Table 1. 99.79% dense samples of dimension 10 mm3 were fabricated using process parameters: power = 90 W, scan speed = 800 mm/s, laser diameter = 80 µm, layer thickness = 25 µm, and hatch spacing = 80 µm.

99.79% dense samples of dimension 10 mm3 were fabricated using process parameters: power = 90 W, scan speed = 800 mm/s, laser diameter = 80 µm, layer thickness = 25 µm, and hatch spacing = 80 µm. Samples were built in the vertical and diagonal direction as shown in Fig. 1. An island scanning strategy with a checkerboard pattern consisting of 5 × 5 mm squares used to build samples is shown in Fig. 1(b). Each island consisted of scan tracks that alternated in opposite directions. Inconel 625 plate (152.4 x 152.4 x 4.76 mm3) was procured from Metalmen, NY, USA. The plate was manufactured per specification AMS5599. Small 1 x 1 x 0.5 cm samples were sectioned from the plate using wire electrical discharge machining (EDM) and used as reference metal.

The plate was manufactured per specification AMS5599. Small 1 x 1 x 0.5 cm samples were sectioned from the plate using wire electrical discharge machining (EDM) and used as reference metal. The chemical composition of wrought IN625 is indicated in Table 1.a) b)Figure 1: a) Inconel 625 samples built in vertical and diagonal orientations with support structures and b) island scan pattern with laser start point (red dot).1802Table 1: Elemental composition as per ASTM standard, L-PBF powder, and wrought IN625 sample used in this study.% wt.NiCrMoFeNb+Ta CoCPSAlTi MnASTMBal20-238-10<5.03.15- 4.15<1<0.1<0.015<0.015 <0.4 <0.4 <0.5L-PBFBal21.599.02.953.550.030.04<0.015<0.0150.110.11 <0.5Wrought Bal22.369.00.723.320.80.030.0090.0050.080.30.182.2.

X-ray diffraction measurementThe XRD technique is the most widely used technique for measuring near surface residual stress as its penetration depth low – on-the-order of 10 µm. Herein, a portable Pulstec μ-X360 X- ray system was employed to carry out residual stress measurements. This system utilizes the novel Cos α method. The samples were irradiated using a beam current of 0.66 mA and voltage of 30 kV. The X-ray beam was directed to be incident on each sample at an angle of 28.8 degrees. The X-ray Manganese (Mn) K-alpha wavelength (λ) used for X-ray diffraction was 2.10 Å. The distance from the specimen and detector to capture the D-S ring image was set to D = 50 mm. The diffraction angle (2was calculated using Bragg’s law, i.e.:nλ = 2dsin𝜃𝜃)(1)where n is an integer and d is the d-spacing or interplanar spacing equal to 1.077 Å in this case.

The diffraction angle (2was calculated using Bragg’s law, i.e.:nλ = 2dsin𝜃𝜃)(1)where n is an integer and d is the d-spacing or interplanar spacing equal to 1.077 Å in this case. Using the above values, the diffraction angle was calculated as 155.03 degrees.𝜃𝜃The μ-X360 X-ray Pulstec system utilizes the Cos α technique to calculate normal residual as presented in Fig 2(a). Keeping the angle of incident stress ( X-rays constant, the relative position of the D-S ring for the L-PBF parts (green ring) from that of the wrought part distortion D-S ring (black ring) was used to calculate strain as presented in Fig 2(b). In this study, the wrought ring position is considered stress free and the L-PBF ring is the stressed ring.

In this study, the wrought ring position is considered stress free and the L-PBF ring is the stressed ring. The (311) family of crystallographic planes were used to create the D-S rings and for verifying the residual stress in as-built and heat-treated samples, as the (311) planes represent bulk macro stresses and are less sensitive to intergranular strain compared to other planes. Since a high level of tensile residual stress is generated at the top surface of the L-PBF samples due to the layer-wise AM process, the top oriented surface of each sample (surface facing upward during printing) was inspected. The equations used to calculate the in-plane stress and shear stress from strains are presented in Eqs. (2)-(5).

EBSD analysisElectron backscattered diffraction (EBSD) was used to study the effects of build orientation on grain orientation, texture, and kernel average misorientation. An analytical FEG-SEM FEI Nova Nano SEM 230 system was used in this study. Samples underwent standard grounding and polishing, followed by vibratory polishing for 2 hours. Similar procedures were performed on wrought IN625 samples to provide reference data. EBSD scans were performed on YZ sections of all samples. Inverse pole figure (IPF) and Kernel average misorientation (KAM) maps were captured in as built vertical, diagonal L-PBF and wrought IN625 samples to estimate local strain levels.3. Results3.1. Residual stressResidual stress magnitudes and nature (tensile or compressive) were determined using the Cos α method along the top surface of the as-built vertically and diagonally printed L-PBF Inconel 625 samples as shown in Fig. 3. As seen from Fig.

3. As seen from Fig. 3, the top surface of both as-built samples is dominated by tensile residual stress with a value of 77 ± 15 MPa and 52 ± 12 MPa, respectively. The vertically printed sample was found to possess higher detrimental tensile residual stress on its top surface compared to the diagonally printed sample surface. Similar results were concluded1804from a study conducted by Pant et al. They measured maximum tensile residual stress in vertically printed IN718 samples [8]. This tensile residual stress is generated from the complex thermal gradients and rapid cooling rates (in the range of 103 K/s to 106 K/s) occurring during the L-PBF process. The intense solidification related to the high-power laser leaving the melt pool and freshly coated powder layers causes built layers under the exposed surface to contract. Peak tensile residual stress value is found in top layers of the built part.

Peak tensile residual stress value is found in top layers of the built part. This is due to the cooling rate provided by the substrate is higher in the underlying layers of the sample. The high cooling rate is increased as the measured point move away from the substrate. Similar results were reported by Shuai et al. where they observed the formation of intense tensile residual stress in as-built L-PBF Inconel 625 samples [15]. In addition, vertically printed samples residual stress measurements had higher standard deviation compared to diagonally printed samples. This implies that the microstructure in the vertically printed sample was more anisotropic.Figure 3: Residual stress measurement using Cos a method in vertically and diagonally printed L-PBF IN625 samples. Error bars represent standard deviation.The D-S rings unique to Cos α method, used to calculate the strain values and ultimately residual stress values in vertically and diagonally printed samples are presented in Fig.

4. The D- S ring, which is recorded on a two-dimensional detector, is taken by the single exposure of X-rays, and this is an important working principle of the Cos α method. In Fig. 4., the cyan ring represents the D-S ring, and the green and yellow rings indicate inner and outer FWHM, respectively. As seen in Fig. 4, the ring distortion from the standard D-S ring position can be used to understand the nature and value of residual stress. The grey lines in Fig. 4 (lines appear as ‘cross-hairs’) indicate the position of a wrought Inconel 625 D-S ring position. Tensile residual stress within a sample shifts the D-S ring position below the grey lines. The presence of compressive residual stress moves the D-S ring position above the grey lines. As seen in Fig. 4., the D-S ring of the vertically built sample is further away from the top grey line compared to the D-S ring of the diagonally built sample.

As seen in Fig. 4., the D-S ring of the vertically built sample is further away from the top grey line compared to the D-S ring of the diagonally built sample. The circularity of the D-S ring recorded in the vertically printed sample is seen to be more distorted compared to the diagonally printed indicating more micro strain present in the microstructure.1805a) b)Figure 4: Debye-Scherrer ring distortion of (a) as-built vertically printed, (b) as-built diagonally printed L-PBF Inconel 625 samples. Cyan, green, and yellow rings indicate D-S ring position, inner and outer FWHM respectively.3.2. EBSDThe difference in residual stress results presented in Fig. 3 can be better understood by performing EBSD to study the crystallography and morphology of grains. The EBSD results of L- PBF IN625 samples were compared with the wrought IN625 sample. Fig. 5 shows the IPF maps of the YZ cross-sections for the as built vertically, diagonally, and wrought samples. As seen in Fig.

Fig. 5 shows the IPF maps of the YZ cross-sections for the as built vertically, diagonally, and wrought samples. As seen in Fig. 5(a), as built sample exhibit texture, with grains oriented in Z- direction. The grains are elongated in Z-direction or build direction, which is inherited from the L-PBF process. A series of columnar grains growing through several layers of powder oriented parallel to the build direction was seen in Fig. 5(a). This is the consequence of directional heat transfer taking place during the solidification stage of printing process. The high cooling rate supplied by the build substrate to every new melted powder layer, helps to adhere with previously melted layer to form elongated grains. This observation is in line with the literatures, where elongated grains along the build direction was reported in L-PBF as built parts [16]. The grains in diagonally printed sample as seen in Fig.

This observation is in line with the literatures, where elongated grains along the build direction was reported in L-PBF as built parts [16]. The grains in diagonally printed sample as seen in Fig. 5(b) were observed to arrange themselves at some angle with respect to build direction. The texture quality in both oriented samples were observed to be similar with minimal variations. The grain morphology in wrought Fig. 5(c) was observed to be equiaxed and oriented in random angles.1806Figure 5: IPF of a) vertical as built, b) diagonal as built and c) wrought IN625 samples. BD stands for build direction. (Color crystalline orientation code is inserted)The kernel average misorientation (KAM) maps presented in Fig. 6 for as built vertically and diagonally oriented, and wrought IN625 samples were used to further analyze the correlation between micro strain caused by dislocations in the microstructure.

6 for as built vertically and diagonally oriented, and wrought IN625 samples were used to further analyze the correlation between micro strain caused by dislocations in the microstructure. The KAM maps show the mean misorientation ratio of a given point and the mean orientation of its neighbors with the same grain. Regions with high KAM values commonly represent regions with plastic strain [17]. The scales enclosed in each KAM maps indicate the level of misorientations. The scales are very low in value which helps address the misorientations caused by dislocations in grains. The micro strain present in the microstructure is highlighted by the green regions. As seen in Fig. 6(a), the micro strain in the as-built vertically printed L-PBF sample is denser compared to diagonally printed sample. indicating presence of high residual stress. The presence of micro strain in the microstructure is coupled with the presence of high dislocation density which is inherited during the L-PBF process.

indicating presence of high residual stress. The presence of micro strain in the microstructure is coupled with the presence of high dislocation density which is inherited during the L-PBF process. Dislocations in diagonally built sample were less dense compared to vertically built IN625 sample in Fig. 3. This confirms the lower residual stress caused by local strain in diagonally printed sample observed in Fig.2. KAM map of wrought IN625 sample is presented in Fig. 6(c) as a benchmark data. The blue circular spots in L-PBF as built samples KAM maps indicate presence of pores arises during the L-PBF build process.1807Figure 6: KAM maps for a) vertically as built, b) diagonally as built, and c) wrought IN625 showing minor misorientations caused by dislocations highlighted in green regions.4. ConclusionsIn this study, the effect of build orientation (vertical and diagonal) on residual stress and microstructure was investigated in L-PBF Inconel 625.

ConclusionsIn this study, the effect of build orientation (vertical and diagonal) on residual stress and microstructure was investigated in L-PBF Inconel 625. A novel Cos α X-ray diffraction method was used to measure surface and sub-surface residual stress. The residual stress results were supported by microstructure analysis via EBSD. The major findings from this study are summarized below:With the same L-PBF system and process parameters, the vertically as built IN625 sample exhibited higher (77 ± 15 MPa) detrimental tensile residual stress compared to the diagonally printed sample (52 ± 12 MPa).The D-S ring captured using the Cos α method was found to be more distorted in the vertically built sample compared to the diagonally built.

This is possibly from stressed grains which diffract the X-ray back to 2D detector in an unorderly path.The inverse pole figure results revealed minimal differences in grain morphology between two build orientations.Kernel average misorientation maps showed that the vertically printed sample inherited dense dislocations which precedes the formation of micro strain and residual stress.The results presented and discussed in this study will enhance the understanding of the effect of build orientation on residual stress in directed energy additively manufactured parts. Residual1808stress can inflict undesirable properties in parts applied in critical industry sectors such as aerospace, biomedical, nuclear, etc. This study can help designers/engineers understand how build orientation effects can be considered when designing and fabricating complex components.AcknowledgementThis material is based upon work supported by the U.S.

Department of Energy’s Office ofNuclear Energy under Award Number DE-NE0008865.DisclaimerThis report was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights. Reference herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof.

Thompson2, Linkan Bian31Center for Advanced Vehicular Systems (CAVS), Mississippi State University, MS 397622Department of Mechanical Engineering, Auburn University, Auburn, AL 368493Department of Industrial and Systems Engineering, Mississippi State University, MS 39762Corresponding author: shamsaei@auburn.eduAbstractIn this study, the effects of build orientation on the mechanical properties and fatigue life of stainless steel (SS) 316L, fabricated using the Laser-Powder Bed Fusion (L-PBF) additive manufacturing (AM) process, were investigated under monotonic tensile and uniaxial strain- controlled fully-reversed (R = -1) cyclic loadings. Tensile tests were conducted at a strain rate of 0.001 s-1, while fatigue tests were performed at strain amplitudes ranging from 0.1% to 0.4% at various frequencies to have a nearly consistent average strain rate in all tests.

The comparison between the tensile properties of additively manufactured and wrought SS 316L revealed that L- PBF specimens exhibited higher yield and ultimate tensile stresses as compared to the wrought specimen. In addition, the elongation to failure of the wrought specimen was similar to that of the horizontally oriented specimen, while it was lower relative to specimens built in vertical and diagonal directions. From the strain-life fatigue analysis, the diagonally oriented L-PBF specimens generally exhibited lower fatigue strength as compared to vertical and horizontal specimens. The fractography analysis revealed three major types of defects to be responsible for the crack initiation and failure.

The fractography analysis revealed three major types of defects to be responsible for the crack initiation and failure. These included (1) voids formed due to lack of fusion between the subsequent layers and entrapped gas, (2) inclusions formed due to the partially melted powder particles, and (3) un- melted powder particles clustered near a void.Keywords: Fatigue, Tensile properties, Fractography, Build orientation, Additivemanufacturing1. IntroductionAdditive manufacturing (AM) processes are gaining attention in both industrial and the academic sectors due to their unique ability to fabricate parts by joining materials layer by layer as opposed to the traditional subtractive fabrication process. As a result, AM processes provide an advantage to produce parts with complex geometry and parts with lower number of assemblies.

As a result, AM processes provide an advantage to produce parts with complex geometry and parts with lower number of assemblies. In general, there are two types of laser-based additive manufacturing techniques for fabricating metallic parts, which are the Laser-Powder Bed Fusion (L-PBF) and the Direct Laser Deposition (DLD) (or Directed Energy Deposition - Laser) processes. In the L-PBF technique, a thin layer of605metallic powder is first deposited on a build plate, which is also known as a substrate. A high intensity laser moving in a pre-defined pattern melts the powder, forming a micro-sized melt pool, which rapidly solidifies once the laser is removed. The solidification process leads to the formation of a solid layer of metal, and this process is repeated for each layer until the entire part is completed.

The solidification process leads to the formation of a solid layer of metal, and this process is repeated for each layer until the entire part is completed. Since the build process involves very high temperatures, the entire process is conducted in an inert environment to minimize oxidation of the fabricated parts.One of the major drawbacks of using the AM process is the susceptibility of its parts to defects, such as inclusions that are formed due to un-melted powder particles, and voids that are created by the lack of fusion (LOF) between the subsequent layers and entrapped gas [1, 2]. Formation of such defects depends on the process parameters such as laser power, hatch spacing, laser pattern, layer thickness, and build orientation [3]. Hence, additively manufactured components may generally exhibit poor fatigue properties when compared to their wrought counterparts, as the defects have significant effects on the fatigue behavior of parts subjected to cyclic loading.

However, such parts often demonstrate comparable or superior mechanical properties as compared to that of wrought material under tensile loading [4]. Further studies on the optimization of L-PBF process parameters are required to fabricate AM parts with better fatigue properties [3].Although various metallic materials may be used for L-PBF, stainless steel (SS) 316L was selected in this study. This particular steel alloy is extensively used in the marine, architectural, and biomedical industries due to its excellent corrosion resistance, weldability, and dynamic stability in wide range of temperature [5]. In biomedical industries, SS 316L is primarily used in orthopedic implants for hip and knee replacements [6], as well as internal fixation of femur bones [7].

In biomedical industries, SS 316L is primarily used in orthopedic implants for hip and knee replacements [6], as well as internal fixation of femur bones [7]. Therefore, as most of the components in the aforementioned applications are subject to cyclic loading throughout their lifetime, it is necessary to perform the fatigue evaluation on SS 316L.Though the applications of SS 316L are popular in biomedical industries, only a small number of research has been performed on mechanical properties of additively manufactured SS 316L [8, 9] and even fewer research have been conducted on their fatigue behavior. Yadollahi et al. [4] investigated the mechanical and microstructural properties of DLD SS 316L. It was found that AM parts possessed higher yield and ultimate stresses relative to their cast and wrought forms. In addition, the elongation to failure was found to be lower for specimens fabricated using the AM process due to defects such as voids and un-melted powder. Sperings et al.

In addition, the elongation to failure was found to be lower for specimens fabricated using the AM process due to defects such as voids and un-melted powder. Sperings et al. [10] investigated the fatigue behavior of two different types of SS (316L and 15-5 precipitation hardening) manufactured using the L-PBF process. It was found that the fatigue life of L-PBF SS 316L has 25% lower fatigue strength at lower strain amplitudes relative to its conventional forms. Hence, further study is required on fatigue performance in order to optimize and select better process parameters for defect-minimized, fatigue resistance fabrication of additively manufactured SS 316L.In this study, the effects of build orientation on the mechanical properties of SS 316L manufactured using the L-PBF process is investigated. First, the experimental setups for conducting tensile and fatigue tests, as well as the fractography analysis are given. The experimental results is then presented and discussed.

First, the experimental setups for conducting tensile and fatigue tests, as well as the fractography analysis are given. The experimental results is then presented and discussed. Finally, some conclusions are made based on the experimental results.6062. Experimental SetupGas atomized SS 316L powder (Renishaw, Inc.), with an average particle size of 25 μm, was utilized for specimen fabrication. Using a scanning electron microscopy (SEM), it was found that the majority of the powder possessed a spherical shape and ranged from 9 to 45 μm in diameter, as shown in Fig. 1(a). The SS 316L rods, with 85 mm length and 8 mm diameter, were fabricated in an argon-purged AM machine (Renishaw AM250) in three different build directions. These are vertically-upward (V), horizontal (H), and diagonal (at an angle of 45° to the substrate (D)), as shown in Figs. 1(b)-(d), respectively. All test specimens were made using default process parameters as presented in Table 1.

1(b)-(d), respectively. All test specimens were made using default process parameters as presented in Table 1. These process parameters, including the laser power, scanning speed, hatching pitch, and layer thickness, were provided by Renishaw Inc. Ten specimens in V and D directions and six specimens in the H direction were fabricated per build using a meander scan pattern. For specimens built in V and D directions, support structures of 5 mm were used only at the base of the cylindrical rod as shown in Figs. 1(b) and (d). Whereas, for specimens built in H direction, support structures were fabricated for every 0.5 mm along the length of the specimen, as shown in Fig. 1(c).After fabrication, L-PBF SS 316L rods were machined using a computer numeric control lathe into round specimens with uniform gage section following the ASTM E606 standard [11] to the dimensions shown in Fig. 2.

2. The gage sections of the machined specimens were then polished using different grit sand papers to remove machining marks and to reduce the surface roughness in the gage section to approximately 1 m.Both monotonic tensile and fatigue tests were conducted at room temperature using a servo hydraulic MTS 858 machine with 25 kN load cells. Axial strain was measured using a MTS mechanical extensometer. The monotonic tensile tests were conducted at a strain rate of 0.001 s-1, first under the strain-controlled loading, followed by the displacement-controlled loading. Since the elongation to failure for the L-PBF SS 316L exceeded the travel distance (4 mm) of the extensometer, the tensile test was switched from strain-control to displacement-control mode to avoid damaging the extensometer. Furthermore, the uniaxial fully-reversed (Rε = -1) strain- controlled fatigue tests were conducted under different strain amplitudes ranging from 0.1% to 0.4% following the ASTM E606 standard [11].

Furthermore, the uniaxial fully-reversed (Rε = -1) strain- controlled fatigue tests were conducted under different strain amplitudes ranging from 0.1% to 0.4% following the ASTM E606 standard [11]. Test frequency for all fatigue tests was varied to maintain a nearly constant average strain rate and to eliminate the effects of strain rate on the fatigue life of investigated L-PBF SS 316L parts. The fractography analysis on the fractured surface of the specimens was conducted using a SEM to investigate failure mechanisms and to determine microstructural factors that are responsible for crack initiation.607(a) (b)(c) (d)Figure 1.(a) SEM image of as-received SS 316L powder along with L-PBF SS 316L cylindrical rods (still attached to substrate) built in (b) vertical, (c) horizontal, and (d) diagonal directions.Table 1.

L-PBF process parameters used for fabricating SS 316L rods.Laser Power (W)Scanning Speed (m/s)Hatching Pitch (mm)Layer Thickness (μm)40010.130608R 3022.57183.580Figure 2.Specimen geometry for both monotonic tensile and fatigue tests (dimensions in mm) based on ASTM E606 standard.3. Experimental Results and DiscussionsMonotonic BehaviorThe engineering stress-strain behavior inherent to the H, D, and V oriented specimens under monotonic tensile loading at a strain rate of 0.001 s-1 is shown in Fig. 3. The comparison of different tensile properties between the L-PBF 316L and wrought SS 316L [4,10,12] was also conducted. These values, including Young’s modulus, yield stress, ultimate tensile stress, and elongation to failure, for wrought SS 316L reported in [4,10,12] are listed in Table 2. From the results, the Young’s modulus for wrought specimen of 200 GPa was observed to be similar to the D specimen and higher relative to the V and H specimens.

From the results, the Young’s modulus for wrought specimen of 200 GPa was observed to be similar to the D specimen and higher relative to the V and H specimens. On the other hand, the yield and ultimate tensile stresses were found to be considerably higher for the investigated L-PBF SS 316L specimens as compared to those reported for wrought SS 316L [4,10]. Among the specimens built in three different directions, the vertically oriented specimen exhibited the greatest elongation to failure (~50% strain), which is approximately 25% higher than that of wrought material. This trend, where the vertically built specimens possessed a higher elongation to failure but lower yield and ultimate tensile stresses as compared to H specimens, also has been reported in [13] for additively-manufactured SS 316L.The SEM images displaying an overall fractured surface for the H, D, and V specimens that failed under monotonic tensile loading are shown in Fig. 4(a)-(c), respectively.

4(a)-(c), respectively. It can be noticed that the gage section of the V specimen, exhibiting the highest elongation to failure, deformed from a circular to an ellipse shape (Fig. 4(c)), while such deformation was not as prominent in specimens fabricated in H (Fig. 4(a)) and D (Fig. 4(b)) directions. Voids due to process-laden gas entrapment and LOF between layers, as shown in Fig. 4, were observed on the facture surfaces of all tensile specimens irrespective to the build orientation.609Figure 3. Monotonic tensile engineering stress-strain curves for L-PBF SS 316L specimens built in horizontal (H), diagonal (D), and vertical (V) directions.Table 2. Tensile properties of wrought SS 316L [4,10,12].Yield Stress (MPa)Ultimate Tensile Stress (MPa)Young’s Modulus (GPa)Elongation (%)255-310535-62320030-40As illustrated in Fig. 3, the elongation to failure was lowest for the horizontally built specimen, while it was comparable for the D and V specimens.

3, the elongation to failure was lowest for the horizontally built specimen, while it was comparable for the D and V specimens. Un-melted powder particles were evident on these fracture surfaces, especially for vertically built specimens. However, these defects apparently did not affect the monotonic tensile behavior of L-PBF SS 316L. The relatively lower elongation to failure observed for H specimens may be related to the longer inter-layer time interval, which resulted in higher cooling rate in these specimens as compared to V and D specimens. This higher cooling rate can in turn lead to finer microstructure in H specimens [4, 12]. Moreover, higher yield and ultimate tensile stresses observed for L-PBF SS 316L specimens (in all three conditions) as compared to the wrought counterpart can be related to the fine microstructure resulting from fast cooling rate during AM fabrication [3, 4, 10, 14].

Further studies on the microstructure, defect statistics, and phase transformation are required to thoroughly understand the effects of build orientation on the tensile behavior of L-PBF SS 316L.610(a) (b)(c)Figure 4.SEM images of overall fracture surfaces of L-PBF SS 316L specimens built in (a) horizontal, (b) diagonal, and (c) vertical directions subjected to monotonic tensile loading.4. Fatigue BehaviorFatigue failure is a common means for mechanical failure, as the majority of components and structures experience cyclic loading throughout their service life. AM induced defects can have significant effects on the fatigue life of AM components by causing stress concentration and providing localized plastic deformation [3]. In general, the fatigue life of a specimen can be categorized, based on the behavior of crack, into three stages; crack initiation, crack propagation, and the final failure.

In general, the fatigue life of a specimen can be categorized, based on the behavior of crack, into three stages; crack initiation, crack propagation, and the final failure. For AM parts, defects have been found to be the main factor responsible for crack initiation [3].Strain-life (εa-Nf) curves, plotted using the experimental data obtained from the uniaxial strain-controlled fully-reversed (Rε = -1) tests for the investigated L-PBF SS 316L specimens, are shown in Fig. 5. In general, among specimens built in three different directions, longer fatigue lives were observed for H specimens, while D specimens exhibited the shortest lives. Build orientation appeared to have less effect on the fatigue life at the highest strain amplitude of 0.4% (i.e. low cycle fatigue). For example, the average life of H specimens tested at 0.1% strain611amplitude is almost three times longer than D specimens, while the average lives of H and D specimens are comparable at 0.4% strain amplitude.

This may be explained by less effects of the crack initiation stage and sensitivity to defects in the short life regime (i.e. higher strain amplitudes) [15].Figure 5.Strain-life curve obtained for L-PBF SS 316L specimens built in horizontal, diagonal, and vertical direction and subjected to uniaxial strain-controlled fully- reversed loading.Fractography analysis was conducted to investigate factors responsible for crack initiation and the failure mechanisms inherent to fatigue specimens. Figures 6-8 display SEM images of fracture surface for the V, D, and H specimens, respectively, which were all subjected to 0.1% strain amplitude at 2.5 Hz. From these figures, three major types of defects were found to be responsible for crack initiation and failure in the L-PBF SS 316L specimens. These are LOF voids between subsequent layers in Fig. 6, inclusions formed due to partially melted powder in Fig. 7, and un-melted powder particles clustered together near voids in Fig.

These are LOF voids between subsequent layers in Fig. 6, inclusions formed due to partially melted powder in Fig. 7, and un-melted powder particles clustered together near voids in Fig. 8.As displayed in Fig. 6, multiple LOF voids, which are typically oriented perpendicular to the build direction, were found to initiate fatigue cracks in the vertically built specimens. These deep LOF defects were observed at the specimen surface as denoted by red circles in Fig. 6(a). The crack coalescences from multiple initiation site, such as those shown in Fig. 6(a) are known to accelerate crack growth rate and may result in shorter fatigue life for V specimens. For the D specimen in Fig. 7, whose fatigue life is generally longer as compared to the V specimen when subjected to an identical cyclic loading condition, the crack was found to initiate from a single larger defect resulting from partially melted powder particles.

This defect appeared to be relatively flat and located close to the surface of the specimen as shown in Fig. 7(b). For the horizontally612built specimen, it was observed that the crack initiated from un-melted powder particles clustered near and inside a LOF void, as illustrated in Fig. 8. This irregular shape LOF void was located slightly away from the surface. In addition, among the fracture surfaces for V, D, and H specimens shown in these figures, it can be noticed that the larger portion of the fracture surface of the H specimen seemed to be under the small crack growth regime as indicated by the smoother appearance, which could result in longer fatigue life for the H specimens. Furthermore, since LOF voids are generally oriented perpendicular to the loading direction in the V specimens and parallel to the loading direction in the H specimens, the effect of LOF voids may be more detrimental on the V specimens as compared to the H specimens [14].

It should be noted that, the SEM analysis presented in this study is only limited to the fatigue specimens subjected to 0.1% strain amplitude (i.e. long life regime). Hence, these observed defects in Figs. 6-8 are not necessarily distinctive for H, D, and V specimens. A comprehensive fractography analysis for all fatigue specimens is required to obtain the influence of build orientation on type and distribution of defects that cause fatigue failure.The SEM images in Figs. 6-8 also indicated that the fatigue life of L-PBF SS 316L was sensitive not only to the types of defects, but also depended on the location and numbers of the defects. Nonetheless, fracture surface analysis alone cannot be utilized to fully explain the fatigue behavior observed in Fig. 5 (shortest fatigue life for D specimens and highest fatigue life for H specimens).

Further studies of microstructural properties, such as grain size and grain orientation, may be required to accurately determine the effect of built orientation on the fatigue life of L-PBF specimens.Figure 6.SEM images of a vertically built L-PBF SS 316L specimen subjected to fully- reversed uniaxial strain-controlled loading at 0.1% strain amplitude and 2.5 Hz showing (a) the overall fracture surface, and (b) void resulting from lack of fusion between subsequent layers that is responsible for the crack initiation.613Figure 7.SEM images of a diagonally built L-PBF SS 316L specimen subjected to fully- reversed uniaxial strain-controlled loading at 0.1% strain amplitude and 2.5 Hz showing (a) the overall fracture surface, and (b) inclusion resulting from partially melted powder that is responsible for crack initiation.Figure 8.SEM images of a horizontally-built L-PBF SS 316L specimen subjected to fully- reversed uniaxial strain-controlled loading at 0.1% strain amplitude and 2.5 Hz showing (a) the overall fracture surface, and (b) inclusion resulting from lack of fusion and un-melted powder that is responsible for crack initiation.ConclusionsIn this study, SS 316L specimens, fabricated using the L-PBF process in three different build orientations (vertical, diagonal (i.e.

45° angle), and horizontal), were subjected to monotonic tensile and uniaxial strain-controlled fully-reversed (R = -1) cyclic loading. Based on the experimental results, the following conclusions can be drawn:6141. All L-PBF SS 316L specimens, irrespective to build direction, exhibited higher yield and ultimate tensile stresses as compared to their wrought counterparts. In addition, when compared to those reported for wrought SS 316L, higher elongation to failure was obtained for vertically and diagonally built specimens.2. Despites the fact that the D specimen possessed better tensile properties (i.e. yield stress, Young’s modulus, and ultimate tensile stress) when compared to specimens oriented in vertical and horizontal directions, shorter fatigue lives were observed for the D specimens, while the H specimens exhibited higher fatigue resistance at low strain amplitudes (i.e. high cycle fatigue).

high cycle fatigue). Minimum effect of build orientation on fatigue life of L-PBF SS 316L specimens was found in the short life regime (i.e. higher strain amplitudes).3. The significant effect of build orientation on the fatigue behavior of L-PBF SS 316L observed at high cycle fatigue can be attributed to greater effects of the crack initiation stage and sensitivity to defects.4. Three major types of defects were observed to be responsible for the crack initiation and failure in L-PBF SS 316L specimens. These defects were voids formed due to the lack of fusion between the subsequent layers, inclusions formed due to partially melted powder particles, and un-melted powder particles clustered near a void.5.

Voids formed due to LOF between layers and their relative orientation to loading direction were observed to play an important role on the fatigue resistance of AM specimens.Based on the results presented in this study, it has been shown that fracture surface analysis alone may not adequately explain the effects of build orientation on the tensile and fatigue behavior of L-PBF SS 316L specimens. Further studies on microstructural properties, such as grain orientation and size may be required. Additional work is also recommended to study the type and distribution of defects formed during fabrication resulting from different combination of process and design parameters, as significant effects of defects were observed on fatigue life of L-PBF SS 316L.AcknowledgementResearch was partially sponsored by the Army Research Laboratory and was accomplished under Cooperative Agreement Number W911NF-15-2-0025.

The views and conclusions contained in this document are those of the authors and should not be interpreted as representing the official policies, either expressed or implied, of the Army Research Laboratory or the U.S. Government. The U.S. Government is authorized to reproduce and distribute reprints for Government purposes notwithstanding any copyright notation herein. This manuscript was prepared while Nima Shamsaei and Scott Thompson were Assistant Professors at Mississippi State University.615References[1] Thompson SM, Bian L, Shamsaei N, Yadollahi A. An overview of direct laser deposition for additive manufacturing; Part I: Transport phenomena, modeling and diagnostics. Addit Manuf 2015;8:36–62.[2] Sterling AJ, Torries B, Shamsaei N, Thompson SM, Seely DW. Fatigue behavior and failure mechanisms of direct laser deposited Ti–6Al–4V. Mater Sci Eng A 2016;655:100–12. [3] Shamsaei N, Yadollahi A, Bian L, Thompson SM.

E EFFECT OF BUILD ORIENTATION ON THE MECHANICAL PROPERTIES IN INKJET 3D PRINTINGJ. Mueller*, K. Shea*Engineering Design and Computing Laboratory, Department of Mechanical and Process Engineering, ETH Zurich, 8092 Zurich, SwitzerlandAbstractIt is known that part orientation plays an important role in 3D printing and especially in inkjet 3D printing, where the layers are more distinct than in other processes. Despite many investigations in this direction, previous research focused mainly on build orientations along the three main axes X, Y and Z. For advanced purposes such as optimization, however, it is important to know what happens in combined alignments between the main axes. The authors hypothesize and show that the transition is not linear and that, despite prior studies, the weakest properties are not found when the parts are aligned along the Z direction.

The authors hypothesize and show that the transition is not linear and that, despite prior studies, the weakest properties are not found when the parts are aligned along the Z direction. The discovered effects can partially be attributed to shear forces in the material, which act between the layers when the parts are not aligned orthogonally to the main axes. To accurately characterize the three-dimensional space, the mixture design method has successfully been introduced to the area of 3D printing and proven to be an efficient tool that can also be used for other processes. With the results of this study, designers and engineers are now able to analyze and predict part properties on a much higher level than before.1. IntroductionAmong the different factors affecting the mechanical properties of fabricated parts, the build orientation is often one of the most influential ones and one of the few factors prevalent in almost all additive manufacturing (AM) processes [1, 2].

The build orientation impacts the accuracy, build time, cost, surface roughness and much more, and its optimization is a fundamental problem in AM which has been tackled by many researchers [3, 4], also from the theoretical side [1, 5, 6]. In many aspects, the effect on the mechanical properties is the most important one as it determines the functionality and not just the visual aspects of a part [7], yet no commercially available software allows for positioning parts on the print table according to the load introduction and direction. One possible reason is that most processes and their effects still aren’t fully understood. In order to get a better understanding, empirical data need to be collected through experiments, which can then be used in models.

In order to get a better understanding, empirical data need to be collected through experiments, which can then be used in models. Extensive research has been undertaken in the area [3, 5, 8], but most of it focused on either orientations along the main axes of the coordinate system, or on continuous orientations within only one plane instead of considering the full three-dimensional space. While in some processes, such as in Fused Deposition Modeling (FDM), the effect of the orientation is relatively small because of indistinct layers [8–10], in other processes such as in inkjet 3D printing, the effect cannot be neglected [11–14]. This is especially true for functional structures such as lattices, where the local load directions of the struts arespatial, even if the global load direction is parallel to a main axis. In order for such structures to perform, they need to be designed and often optimized [15, 16], which is only possible when accurate material data is available.

In order for such structures to perform, they need to be designed and often optimized [15, 16], which is only possible when accurate material data is available. Gathering accurate material data for all possible build orientations, however, can be costly. A trade-off needs to be found for the required accuracy and the experimental cost. Hence, the mixture design, a statistical tool commonly used in chemistry, will be adapted and introduced into the field. It is tested for its applicability and robustness before conducting the main tests for the most common load cases compression and tension. The work is based on the inkjet 3D printing process, and the trends of the results and models are valid also for similar processes.In the background section we discuss the relevant literature and put it into perspective with other AM processes. We then describe the materials and methods, and present the results in section four.

We then describe the materials and methods, and present the results in section four. A discussion of the results is provided and conclusions are drawn in the last section of the report.2. BackgroundAdditive manufacturing or 3D printing is defined as a layer based manufacturing process [17]. This statement generally holds true, but, due to the large amount of different printing processes, sometimes the layers are not very distinct. For example, in Selective Laser Sintering (SLS) and Fused Deposition Modelling (FDM), the material is cured voxel by voxel, and not line by line or layer by layer. Hence, the bonding strength between two layers is identical to the bonding between two printed lines within a layer, with only a difference in time. Looking at the global properties, the printing or curing path can be designed such that almost isotropic parts are created.

Looking at the global properties, the printing or curing path can be designed such that almost isotropic parts are created. Isotropy, however, is not always the goal as there is often only one load direction required for which the mechanical properties can be optimized. In those cases, the anisotropy is enforced. Less freedom is available in inkjet 3D printing, where the variances in terms of the mechanical properties sometimes reach 30%, which has been investigated on different printer types: Objet30 [19], Objet Eden 260 [18] and Objet Eden 330 [20–22]. The Objet350 Connex, which was investigated by Barclift and Williams [11], is closest to the printer used in this work, the Stratasys Objet500 Connex3. On the Connex3, the effect of the build orientation on the mechanical properties was investigated by Mueller et al. [13, 14]. However, all of these investigations have in common that they only look at build orientations along the three main axes.

[13, 14]. However, all of these investigations have in common that they only look at build orientations along the three main axes. In this paper, the transition between the main axes is investigated for all possible build orientations. Since the Connex3 is one of the most advanced commercially available multi material printers [23], it can be considered to be representative for the current state-of-the-art in commercial inkjet 3D printing.In comparison to inkjet 3D printing, the variations in FDM can range from about 9% in the strength [9], over 11% in both modulus and strength [8] to 28% in strength [10]. The large differences can mostly be attributed to the different print settings. For SLS, variations ranging from 7% to 47% are reported [24]. Note that in both of these processes the anisotropy is intended and can be drastically reduced by optimized alignment of the print or cure path, which is not the case in commercial inkjet 3D printing.

In inkjet 3D printing, a print head with multiple, linearly aligned nozzles moves back and forth positioning the liquid material before curing it almost homogeneously through the attached ultra-violet (UV) lights. Hence, the layer itself is built in amuch shorter time compared to other processes. This yields a smaller anisotropy within the layer, but a larger one between the layers [13, 14], creating a structure with transversally isotropic properties. Roughly speaking it can be said that the smaller the ratio𝜈𝑓 =𝑡𝑖𝑚𝑒 𝑡𝑜 𝑝𝑟𝑖𝑛𝑡 𝑜𝑛𝑒 𝑙𝑎𝑦𝑒𝑟 𝑡𝑖𝑚𝑒 𝑡𝑜 𝑠ℎ𝑖𝑓𝑡 𝑏𝑒𝑡𝑤𝑒𝑒𝑛 𝑙𝑎𝑦𝑒𝑟𝑠the higher the minimum anisotropy. This, however, cannot be generalized and needs to be investigated in detail for each individual process. 𝜈𝑓 is also an indicator for a number of other effects, such as the printing time [25–28].

This, however, cannot be generalized and needs to be investigated in detail for each individual process. 𝜈𝑓 is also an indicator for a number of other effects, such as the printing time [25–28]. For a process with a small 𝜈𝑓, such as the inkjet 3D printing process, the print time drastically increases when the build orientation changes, whereas the change is much smaller for processes with a high 𝜈𝑓, like FDM [29, 30]. Since the build orientation is one of the early steps in the AM cycle, it has additional importance since it subsequently influences many following steps such as the slicing, support generation, tool path definition, fabrication, part removal and cleaning [7].3. Materials and MethodsThe experimental procedure is described based on the experimental design, printing and testing.3.1 Experimental designA response surface is applied to multiple measurement points which are analyzed experimentally.

More specifically, an unaugmented mixture design [31] with three components is chosen that allows to define the proportions of the coordinate system, i.e. the X, Y and Z axes, as contributions to a blend. The axes are defined as components or factors of the design, as shown in Figure 1. The dots are the levels or settings of the factors. Then, a detailed in-plane measurement is conducted for the tension’s Young’s modulus in the XZ plane to investigate the suitability of the chosen experimental design and its settings, indicated in blue. Angles ranging from 0° to 90° are measured in 10° steps on 5 specimens for each point and compared to second and third polynomial degree models, which represent the simplex centroid (left) and simplex lattice degree 3 (right) models. The results are analyzed to find the best trade-off between accuracy and experimental cost, and the more suitable design is chosen for the main tests accordingly.Figure 1: Unaugmented mixture designs.

The black dots indicate the set measurements in relation to the axes and the blue dots indicate the in-plane measurements of the XZ plane.3.2. PrintingAll test-specimens are printed on a Stratasys Objet500 Connex3 printer under air- conditioned temperatures. Each test run is printed in a separate tray and the print head is cleaned before each run to prevent an influence through blocked nozzles [13, 14]. VeroWhitePlus (RGD835) model material from one batch is used in combination with SUP705 support material. To ensure a uniform surface finish and to prevent over-curing [19], the printers’ mat option is used, with which the printer covers all surfaces in support material rather than adding it exclusively to overhanging features. Due to the slenderness, the specimens parallel to the Z axis are printed in an additional support structure made from model material. All parts are cleaned of support material with a Krumm RK 5 XL VA water-jet.3.3.

All parts are cleaned of support material with a Krumm RK 5 XL VA water-jet.3.3. Mechanical testingThe tests are conducted on an Instron ElectroPuls E3000 universal testing machine. For the compression tests, the settings were chosen according to the ASTM D695-10 [32], i.e. the standard test specimen in the shape of a right cylinder with ’preferred dimensions’ of 12.7 mm by 12.7 mm by 25.4 mm and a standard speed of testing of 1.3 mm/min. For the tensile properties, the ASTM D638-10 [33] standard is used. In order to fit the vertical build orientation on the tray, specimen geometry of type IV is chosen in combination with a constant testing speed of 50 mm/min. At least five replicates were tested for each setting and shuffled in print and test order to minimize random effects. As part of the mixture design, an Analysis of Variance (ANOVA) is conducted for each test, which provides information on the statistical significance of the results of different comparison groups.

As part of the mixture design, an Analysis of Variance (ANOVA) is conducted for each test, which provides information on the statistical significance of the results of different comparison groups. A critical p-value of pcrit=0.05 is chosen, indicating that effects with p < 5% are significant.4. ResultsFollowing, the empirical results of the tests are presented, starting with the in-plane tests. Then, the results of the mixture design tests are presented, followed by the application of the regression model, which yields the response surfaces.4.1. In-plane testsThe results of the in-plane measurements of the XZ plane are shown for Young’s modulus in Figure 2 with a black line. Curves for the simplex centroid and simplex lattice designs are added in red and green, respectively. Both curves are fitted to the measurement data in the start and end points, and at 45° for the centroid and at 30° and 60° for the simplex lattice.

Both curves are fitted to the measurement data in the start and end points, and at 45° for the centroid and at 30° and 60° for the simplex lattice. In blue, a linear interpolation between the start and end points is shown. All lines start at 2712 MPa. Between 0° and 45°, the interpolated measurement curve is well approximated by the simplex centroid and more so by simplex lattice, and all of those curves meet at 45° at about 2500 MPa. While the measurement and simplex lattice curves then further drop until they reach about 2472 MPa, the approximate minimum is reached at 45° for the simplex centroid. In this second half of the graph, the simplex lattice better approximates the measurements and at 90° all curves meet at 2600 MPa.4.2. Mixture design testsCompression. A summary of the compression test results shown in Table 1 and the final results of the mixture design, after applying the regression model, are shown in Figures 3a and 3c.

A summary of the compression test results shown in Table 1 and the final results of the mixture design, after applying the regression model, are shown in Figures 3a and 3c. The shapes of the fitted curves are similar for both mechanical properties and the build orientation along Y is the strongest with E = 2396 MPa and US = 87 MPa. Among the main axes, the lowest values are Ec,x = 2272 MPa and USc,z = 85 MPa, with the remaining values laying in between at Ec,z = 2358 MPa and USc,x = 86 MPa. The indices c and t stand for compression and tension, respectively, and are followed by the build orientation.Figure 2: In-plane measurements (XZ) compared to linear and polynomial interpolations.The properties in between two axes are always weaker than in the main directions and reach their minimum at approximately half way at 45°: Ec,xy = 2177.63 MPa, USc,xy = 84 MPa, Ec,xz = 2209 MPa, USc,xz = 84 MPa, Ec,yz = 2171 MPa and USc,yz = 85 MPa.

The weakest properties are found in between the three axes, close to the 0.5x + 0.5y + 0.5z orientation (Ec,xyz = 2135 MPa, USc,xyz = 84 MPa).Tension. The tension test results are shown in Table 2. The final tension test results after applying the regression model are shown in Figures 3b and 3d. For the Young’s modulus, the shape is similar to the compression results: The peaks are at the main axes with lower values laying in between.The values at X and Y are about the same (Et,x = 2765 MPa, Et,y = 2778 MPa), with the Z axis being considerably below (Et,z = 2689 MPa). The intermediate values are always lower and reach their minima at about half way (Et,xy = 2684 MPa, Et,xz = 2530 MPa, Et,yz = 2579 MPa). The lowest value found is, again, near the center at 0.5x + 0.5y + 0.5z (Et,xyz = 2539 MPa). For the ultimate strength, the difference between the X and Y axes is marginal (USt,x = 62 MPa, USt,y = 66 MPa, USt,xy = 64 MPa).

For the ultimate strength, the difference between the X and Y axes is marginal (USt,x = 62 MPa, USt,y = 66 MPa, USt,xy = 64 MPa). This is true for all levels of Z, which has the lowest value at its main direction (USt,z = 41 MPa). As opposed to the compression results, the intermediate values between the X and Z, and Y and Z axes show a convex trend (USt,xz = 57 MPa, USt,yz = 58 MPa).Table 1: Compression test results for different print orientations. The numbers in the bracelets depict the standard deviation. E [MPa] Y X 0 2351 (151) 0 1/2 1/2 2200 (196) 0 0 2389 (249) 1 1/3 1/3 1/3 2070 (120) 1/2 2237 (227) 1/2 0 2206 (212) 1/2 1/2 0 2265 (180) 0 0 1Z 1UTS [MPa] 85.12 (1.02) 84.99 (0.53) 87.40 (0.80) 83.18 (0.78) 84.14 (0.73) 84.09 (0.62) 86.10 (0.89)0Table 2: Tension test results for different print orientations. The numbers in the bracelets depict the standard deviation.

Part of the mixture design is an ANOVA, which provides information on the statistical relevance of the tests in terms of a p-value. The results for the tests and source terms are shown in Table 3. Since the linear terms are included in the quadratic terms, the quadratic terms are always of higher relevance. This is the case for the p-Values of Ec: With p=0.396, the linear term is far above the critical value pcrit=0.05. However, the quadratic terms (p=0.008) show that there exists a strong significance. This is especially true for the X*Y (p=0.041) and Y*Z combination (p=0.008). The notation X*Y indicates interaction effects between the named factors. The value for X*Z (p=0.160) is slightly above the critical value, but looking at the graph (Figure 3a) confirms the likelihood that the effect is also significant.

For USt, the linear terms show a strong significance (p<0.001), and the overall quadratic terms are slightly above the threshold (p=0.097). The X*Y term does not show a significance, which is in agreement with its graph (Figure 3d). The X*Z and Y*Z significances (p=0.039 and p=0.133) are, however, close to the threshold and the graph shows indeed a strong decline towards Z.5. DiscussionFirst, the design model decision is established based on the in-plane tests. Then, the main mixture design results tests are discussed and put into relation with a traditional process, injection molding.5.1. In-plane testsAlthough previous research did not specifically interpolate between the measurements, it is assumed that the whole width of the anisotropy is covered by inspecting build orientations only along the main axes.

It is recognized that this approach is not sufficient when looking at the results of the mixture designs, but even more so when looking at the detailed in-plane measurements which are also compared to the linear interpolation (Figure 2). In contrast, an additional interpolation point in combination with a polynomial fit (degree 2), similar to the mixture design approach, drastically reduces the deviation error. Two additional points (polynomial degree 3) further increase the fitted accuracy, but add at least three measurement points per replication (Figure 1) or, in other words, 43% experimental cost. For this reason and due to the large number of tests conducted in this work, it has been decided to proceed with the simplex centroid design.5.2. Mixture design testsThe effect of the layers in inkjet 3D printing can be described, in a simplified way, as polymer sheets, i.e. laminae, glued on top of each other to build adhesive joints.

Mixture design testsThe effect of the layers in inkjet 3D printing can be described, in a simplified way, as polymer sheets, i.e. laminae, glued on top of each other to build adhesive joints. This is due to the fact that the liquid material is cured layer-wise, hence the intralaminar bonds are stronger than the interlaminar bonds. A similar phenomenon in composite materials is termed shear extension coupling: For out-of-plane alignments, i.e. when the normal stresses are not aligned with the printer’s axes, the normal stresses cause shear strains and shear stresses cause extensional strains [34].Compression. In in-plane loads, i.e. along X and Y, and in loads orthogonal to the plane, i.e. along Z, the loads are not straining the bonds directly. For the angles in between, shear forces are created which reach their maximum at about 45°, an effect also seen in the experimental results.

along Z, the loads are not straining the bonds directly. For the angles in between, shear forces are created which reach their maximum at about 45°, an effect also seen in the experimental results. As these forces directly affect the bonds, the mechanical properties are decreased (Figures 3a, 3c).Tension. The same is valid for the tension, with the difference that in the Z orientation, the load also attacks the bonds between the layers. For the stiffness, the effect is visible, but weak, which is probably due to the fact that, as long as the bond lasts, the stiffness is not affected. The strength, however, is drastically reduced towards Z. This goes along with the recommendations for preferred load cases in adhesive bonds [35], where shear is preferred to normal loads.

The strength, however, is drastically reduced towards Z. This goes along with the recommendations for preferred load cases in adhesive bonds [35], where shear is preferred to normal loads. That also has to do with the fact that perfect orthogonality cannot be reached in practice [35].On a more general level, the mechanical properties of inkjet 3D printed VeroWhitePlus material found are similar to other studies: Et = 2.7 to 2.8 GPa and USt = 41 to 69 MPa [13, 14]. Stankovic et al. [15, 16] also investigated the Poisson’s ratio and found values of 𝜈 = 1.17 to 1.18 g/cm3. In comparison to other polymers, a strong similarity with Nylons is found, which has E =1.8 to 3.5 GPa, 𝜎𝑦 = 49 to 87 MPa, 𝜎𝑇𝑆 = 100 MPa and 𝜈 = 1.1 to 1.2 g/cm3 [36]. This information allows to make educated guesses of unknown properties of the used material, such as shear moduli, which are important in numerical studies.

This information allows to make educated guesses of unknown properties of the used material, such as shear moduli, which are important in numerical studies. It is also common in other, more traditional manufacturing processes that the resulting parts have mechanical anisotropies [37, 38]. For example, thermoplastic materials have different properties in the flow direction compared to the cross-flow direction, which is caused when the shear stresses in the flow orient the polymer chains, leading to higher strength. There are also other properties that induce anisotropy, such as injection speed, melt and mold temperature, shrinkage and wall thickness [38]. In non-optimized cases, the variations in the mechanical properties can exceed 25% [37], which makes it comparable to the inkjet 3D-printing process. An advantage of the anisotropy of AM compared injection molding, however, is that the direction of the anisotropy can often be actively controlled with less restrictions.

An advantage of the anisotropy of AM compared injection molding, however, is that the direction of the anisotropy can often be actively controlled with less restrictions. Since the software on the printers do not have information on the load introduction and direction, other objectives are used for the optimization of the print orientation; usually the cost in terms of required support material and printing time. It is assumed, however, that in the future, software developers will include a feature to find the optimal build orientation for a given load case, for example [39].6.

It is assumed, however, that in the future, software developers will include a feature to find the optimal build orientation for a given load case, for example [39].6. ConclusionsFirst, for the first time, a mixture design was used to measure the effect of build orientation and, although having its roots in different areas, has proven to be an accurate and efficient tool to analyze the mechanical properties in a three-dimensional space.Second, the findings of this study indicate that, when considering all possible build orientations, comparable properties to established processes such as injection molding can be reached in terms of anisotropy.Third, the anisotropy in inkjet 3D printed parts has been investigated on a state-of-the-art printer to a high level of accuracy.

The findings will inform designers, enable better predictions and create stronger parts by taking into account the dominating effects of shear stresses when the load directions are not aligned parallel to the main axes.AcknowledgementsThe research is supported by the ETH Zurich, Seed Project SP-MaP 02-14, ”AdditiveManufacturing of Complex-Shaped Parts with Locally Tunable Materials”.References[1] P. Kulkarni, A. Marsan, D. Dutta, A review of process planning techniques in layered manufacturing, Rapid Prototyping Journal 6 (2000) 18–35. [2] R. Hague, S. Mansour, N. Saleh, Material and design considerations for rapid manufacturing, International Journal of Production Research 42 (2004) 4691–4708. [3] P. Alexander, S. Allen, D. Dutta, Part orientation and build cost determination in layered manufacturing, Computer- Aided Design 30 (1998) 343–356. [4] D. Frank, G.

H. Hernandez1,2 L. Martinez1,2 M. I. Lopez,1,2 R. B. Wicker2,3 and S. Collins4 1Department of Metallurgical and Materials Engineering The University of Texas at El Paso, El Paso, TX 79968 USA 2W. M. Keck Center for 3-D Innovation The University of Texas at El Paso, El Paso, TX 79968 USA 3Department of Mechanical Engineering The University of Texas at El Paso, El Paso, TX 79968 USA 4Arcam AB, Mölndal, Sweden Abstract In this study, involving additive manufacturing (AM) using electron beam melting (EBM), we have examined build defects which result from beam tripping, porosities (including unmelted or unsintered zones) due to excursions from optimal build parameters (especially variations in melt scan beam current and scan speed), and gas bubbles trapped in atomized Ti-6Al-4V starting powder as well as recycled powder, and retained in the build.

At optimized build conditions we have also examined microstructure-mechanical property (hardness, tensile strength, and elongation) variations for multiple rake building and multiple melt scans using optical metallography and scanning and transmission electron microscopy (SEM and TEM). These build variances cause cooling rate variances which promote α-phase growth and variations in dislocation density, as well as α-to-α' (martensite) phase changes, all of which produce some degree of mechanical property variations. These features (especially α-to-α' phase changes) are notable on comparing solid builds in comparison with a variety of mesh arrays where strut dimension and build-element complexities alter the cooling rates in a significant way. We illustrate these microstructure variations with corresponding variations in microindentation hardness measurements made directly on fine mesh (strut) structures.

We illustrate these microstructure variations with corresponding variations in microindentation hardness measurements made directly on fine mesh (strut) structures. Finally, we have examined Ti-6Al-4V powder chemistries and solid build chemistries which for single-pass melt scans at optimized build conditions are shown to be relatively constant up to 40 cycles of powder reuse with the exception of Al content which was reduced by 10 to 15% in solid builds at optimized conditions. However, Al loss in solid builds approached 25% for multiple (2 and 3) melt scans, while no changes in Ar gas-bubble density were observed with changes in α-phase (grain) width which increased from 3 µm for a single melt scan to 4.5 and 6 µm for 2 and 3 melt scans, respectively.

Corresponding Rockwell C-scale (HRC) hardness varied from 37, 36, and 35, respectively; with ultimate tensile strengths exceeding 1.2 GPa at elongations of 12% or higher for this melt scan sequence.Introduction The fabrication of complex, multifunctional, monolithic components is a particular advantage of solid freeform fabrication (SFF). For example, while biomedical devices can be machined from forged billets and proximal or total surface porosity developed by subsequent spray or sinter processes, it is impossible to join foam or mesh components of various kinds as an integral part of a monolithic device. Furthermore, metal mesh or foam structures are impossible to produce as an integral part of the conventional machining process, or any related manufacturing process except for foaming protocols in the melt, which do not work for Ti-6Al-4V and other high-temperature, reactive metals and alloys.

Similarly, mesh or foam arrays for impact energymanagement in aerospace and automotive applications are not conventionally tractable as well as numerous, related product configurations for thermal management (such as heat exchangers), sound adsorption and vibration suppression structures, and the like. However, additive manufacturing (AM) using laser or electron beams can build any complex, monolithic component through computer-controlled or computer-aided design (CAD) driven self-assembly by melting precursor powder layers.

Unfortunately, these processes are not well certified in terms of manufacturing standards which are well documented for wrought or cast precursor manufacturing for Ti-6Al-4V in particular, and related alloy systems, although there has been some recent progress for EBM produced components (Wooten, 2008; Gaytan et al., 2009).In this research program, we have been concerned with optimizing the EBM build process and understanding microstructural and mechanical variations which accompany variations in build parameters and build geometries using Ti-6Al-4V precursor powder.Methods and ProceduresFigure 1 illustrates the ARCAM (A2) EBM system and AM process (Gaytan et al., 2009). Similar in concept to other electron optical devices, an electron beam is generated in an electron gun ((1) in Fig.

Similar in concept to other electron optical devices, an electron beam is generated in an electron gun ((1) in Fig. 1.) and projected through a focusing lens (2) and subsequently rastered over the building layers of powder (3), first involving multiple scans (~11) at high rate and high current (~15,000 mm/s @ 30mA) to preheat the powder layer to 760º C, leveling to ~640º C, followed by an optimized melt scan at ~400 mm/s @ 6 mA. Varying either the scan rate or current as well as the imposition of multiple melt scans can have very noticeable effects on the efficiency of the build as well as the properties of the build; microstructure and mechanical properties. The mean or average powder size illustrated in the insert in Fig. 1 was roughly 30 µm. Consequently each nominal 100 µm thick melted layer corresponds to several (≥ 3) mean powder particle layers. The powder is gravity fed from cassettes ((4) in Fig. 1) and raked (5) across each melt layer.

The powder is gravity fed from cassettes ((4) in Fig. 1) and raked (5) across each melt layer. Multiple rake passes for each layer may also alter the build properties by more effectively filling layer interstertices with smaller powder particles. Consequently, beam scan and current variances, melt scan variances and rake variances, can all have an effect on the build product and build efficiency. These were explored in this study.The powder used in this study was Ti-6Al-4V ELI powder having a nominal composition of 6.04 percent Al, 4.05 percent V, 0.013 percent C, 0.07 percent Fe, 0.13 percent O, and <0.005 percent N and H; the balance Ti (in weight percent). The average powder particle diameter was ~30 µm as noted. We have analyzed the precursor powder in the SEM (a Hitachi S4800 field- emission SEM) utilizing an energy dispersive (X-ray) spectrometer (EDS) as well as recycled powder, and the builds themselves to observe elemental variations, especially volatilization of Al.

While oxygen is detectable by EDS, the resolution is ≥ 0.5% (weight).Microstructures of various builds, including mesh arrays and related geometries were examined by optical metallography by mounting selected sections or coupons which were ground, polished, and etched; using a solution consisting of 100 mL H2O, 2.5 mL HF, and 5 mL HNO3. Polished and etched specimens were also tested by microindentation hardness testing using a Vickers indenter at 25 to 100 gf loads (HV). Solid, fully dense monoliths were also tested by Rockwell C-scale hardness indentation, and selected samples were machined into tensile specimens and tested using an Instron 500 R tensile machine at a strain rate of 3 x 10-3 s-1,Figure 1. EBM schematic and SEM view of precursor powder (insert). Numbers denote the system components: (1) Electron gun, (2) Beam focus lens, (3) Beam deflection coils, (4) Powder cassettes, (5) Powder layer rake, (6) Build product, (7) Build table. From Gaytan, et al.

From Gaytan, et al. (2009).at room temperature (~20ºC). HV and HRC values are nearly corresponding except for an order of magnitude or decimal place difference: HV (xxx) ~ HRC (xx.x).We utilized techniques previously described in detail (Murr et al., 2009a) to observe microstructures in builds by transmission electron microscopy. However, since we have recently demonstrated that mesh arrays and foams of Ti-6Al-4V fabricated by EBM exhibit variations in microstructure as a consequence of different cooling rates for small geometrical components such as mesh struts (Murr et al., 2009b), we developed a novel build concept to examine these issues by TEM. Since it is not possible to prepare electron transparent thin sections for TEM from 1 mm or smaller struts, we developed a “test box” concept illustrated in Fig. 2. In this concept, l cm2 window arrays having various thicknesses (ranging from 0.5 mm to 2 mm) are created along with 1 cm3 solid sections for comparison (Fig. 2(b)).

2. In this concept, l cm2 window arrays having various thicknesses (ranging from 0.5 mm to 2 mm) are created along with 1 cm3 solid sections for comparison (Fig. 2(b)). Standard 3 mm discs for TEM specimen preparation were then punched from these “windows” after grinding to ~0.2 mm thickness, including coupons cut from the solid test blocks (1 cm3) in Fig. 2. These punched discs were electropolished as described previously (Murr et al., 2009b), and examined in a Hitachi H-8000 analytical TEM operating at 200 kV accelerating potential, employing a goniometer-tilt stage. Similar “windows” in Fig. 2 were also mounted, polished, and etched for optical metallography in order that the macrostructure observed by optical metallography could be correlated with the TEM images. The hardness of “window” sections cut and mounted from Fig. 2(b) was also measured using a Vickers indenter (HV) as described previously.

The hardness of “window” sections cut and mounted from Fig. 2(b) was also measured using a Vickers indenter (HV) as described previously. Several variations of this concept were pursued in order to achieve requisite cooling and cooling-rate variations.Results and DiscussionBuild Defects and Related Issues There are notable build defects which can occur by improper layer filling, beam tripping which can involve momentary variations in beam energy or beam blanking. These involve poorly melted (or sintered) regions which can be considered porosity zones, and examples are illustrated in Figure 3. In addition, variations in optimized build parameters such as melt scan beam current or scan rate can also have deleterious effects on build integrity or build defects. Figure 4 illustrates this feature in the extreme by changing the melt scan beam current from the optimum of 6 mA to 1 mA. This creates an extremely porous and poorly consolidated build.

Figure 4 illustrates this feature in the extreme by changing the melt scan beam current from the optimum of 6 mA to 1 mA. This creates an extremely porous and poorly consolidated build. Note the poorly sintered zone represented in Figure 4(b). This example also illustrates that it is not possible to build porous structures with any real integrity by simply changing melt scan rate or beam current.We have examined the prospects for defects in builds depending upon whether the powder is raked one or more times for the creation of each layer. No statistically significant differences were observed on comparing a single rake pass with two (or three) rake passes, which is the preferred procedure for optimized builds. The reasoning is of course the fact that layer filling would be somewhat more complete for multiple rake passes.

The reasoning is of course the fact that layer filling would be somewhat more complete for multiple rake passes. However, the Ti-6Al-4A powder illustrated in the insert in Figure 1 shows smaller particles sintered to larger ones, producing a bimodal particle size distribution as illustrated previously by Murr et al. (2009a). This may assure more consistent layer filling in contrast to unimodal particle size distributions.Because the atomization process for manufacturing uniform, round Ti-6Al-4V particles traps argon gas bubbles, these bubbles are carried into the melt and retained in the solidified build asFigure 2. CAD model (a) and actual Ti-6Al-4V build prototype (b) for build “window” array concept. The windows or similar array components as shown at “A” in (a) or “windows” have varying thicknesses to produce varying cooling rates. These array components permit convenient characterization by optical microscopy and TEM.

These array components permit convenient characterization by optical microscopy and TEM. TEM requires 3 mm discs to be punched from the thinnest component (~ 0.2 mm) while thicker features are ground and polished to this thickness. The surface features in (b) are exaggerated by spectral reflectance.Figure 3. Examples of build defects. (a) Unmelted and consolidated zone. (b) is similar to (a) with gas void (exposed bubble) at arrow. From Gaytan, et al. (2009).Figure 4. Ti-6Al-4V build comparisons. (a) Shows the optimized build parameters to produce fully dense monolith while (b) shows a reduction in melt-beam current. From Gaytan, et al., (2009).illustrated in Figure 5 which shows retained argon bubbles which can range in size from <10 µm to >50 µm in diameter. The argon gas is apparently not under any significant pressure when trapped, and the high temperature of the melt (~1670º C) and the high surface tension of Ti-6Al- 4V does not permit these bubbles to leave the melt.

Figure 6 illustrates several typical examples of bubbles breaking out of the precursor powder particles, especially in the SEM view shown in Figure 6b.As noted, the high melt temperature and surface tension for Ti-6Al-4V does not allow the argon bubbles to escape. This feature is illustrated in a somewhat exaggerated form for reconstructed, 3D sections representing a range of melt scans: 1, 2, and 3 as shown in Figure 7a, b, and c, respectively. It is apparent that while the bubble distribution is not significantly altered with multiple melt passes, the microstructure is changed because of the variations in heat content and cooling rate.

It is apparent that while the bubble distribution is not significantly altered with multiple melt passes, the microstructure is changed because of the variations in heat content and cooling rate. This is illustrated by the systematic increase in the acicular α-plate dimensions with the number of melt scans in Figure 7 a to c, respectively.We have also recently shown that Ti-6Al-4V powders can be continuously recycled with no significant variation in elemental concentrations where Al concentrations averaged over hundreds of individual powder particles ranged from 4.3 to 7.3 weight percent Al and 2.9 to 3.8 weight percent V for the original (precursor) powder as illustrated for example in Figure 8 (Gaytan, et al., 2009).

However, there was a notable variation in Al content for single melt pass monoliths versus three-melt pass monoliths: ranging from 10 to 15% for single and double melt pass (fully dense)monoliths (Figure 7a and b) in contrast to roughly 15 to 25% reduction in average Al content for triple (3) pass melt scans (Figure 7c). There was no systematic V loss between these melt scan/pass monoliths. Thermal Management Microstructures and Mechanical Properties of Monolithic (Solid) and Complex Mesh Builds It is apparent from Figure 7 that heat content and thermal management (cooling rate, etc.) can have a significant effect on the microstructure and associated mechanical behavior – such as residual hardness – for monolithic, solid products built by EBM. It has also been observed in this context that because heat content varies from the bottom to the top of a solid build-even for simple cylinders the microstructure and associated hardness will vary (Murr, et al. 2009a).

2009a). Additionally, Murr, et al. (2009c) have recently demonstrated that complex mesh or open cellular structures can have variations in acicular α-phase platelets and α'-phase (hcp) platelets which produce a finer Widmanstätten structure, having higher residual hardness. Correspondingly, Murr, et al. (2009b) have shown that α' (martensite) phase platelets contribute to strengthening, but result in lower ductility (as measured by tensile elongation).Figure 9 shows a CAD model and examples of actual builds for a 3D-reticulated mesh structure exhibiting non-isotropic hexahedral and octahedral (Figure 9c and d) geometrical arrays for a build element referred to as “dode-thin” from Materialize software (Murr et al. 2009c). These mesh arrays are composed of varying size struts or truss rods having diameters ranging from 0.5 to 1.5 mm, and with varying spacings which give rise to variations in residual, product porosities and densities as well as thermal or heat flow variations.

Figure 10 shows an SEM view of these interconnected strut arrays along with a typical optical metallography view of the associated α' (martensite) platelets which can be contrasted with the variations in α-phase (hcp) acicular grains illustrated in Figure 7. Figure 11 shows for comparison the enlarged α-phase,Figure 5. Examples of gas (void) bubbles in solid monoliths. (a) and (b) are optical metallograph images while (c) and (d) show SEM images.Figure 6. Gas bubbles in Ti-6Al-4V powders (typical for precursor/starting powder and recycled powder). (a) SEM metallographic view of mounted, ground, polished, and etched powder particle showing interior void. (b) SEM view of gas bubble breaking the powder particle surface. From Gaytan, et al. (2009).Figure 7. Series of 3D reconstructions of optical metallograph views for solid monoliths built with optimum melt scan parameters, but with (a) one melt scan, (b) two melt scans, (c) three melt scans.

In Figure 13a, the cross-sections for the foam structure are roughly 1 mm in dimension while some structural units vary from 1 mm wide to 3 mm length. Note the layering implicit in the magnified SEM view shown in Figure 13b.It might be expected that like the interconnected, small strut structures composing the mesh arrays illustrated in Figures 9 and 10, the cellular structures in Figure 12 and 13 would be expected to exhibit rapid cooling and some propensity of α'-martensite platelets. This feature is illustrated in Figure 14f in contrast to the α-phase platelet variation for increasing melt scans.

This feature is illustrated in Figure 14f in contrast to the α-phase platelet variation for increasing melt scans. (Figure 14a to c) and the variation of α-phase acicular grain structure from the top to the bottom of a simple, single melt scan cylindrical build shown respectively in Figure 14d and e.Figure 15 compares the TEM microstructures which specifically correspond to the bottom, harder α-phase structure shown in Figure 14e (Figure 15a and b) and the 3-melt scan microstructures shown in Figures 7c, 11a, and 14c. The corresponding hardness difference from Figure 15 a and b to Figures 15 c and d represents a notable difference in dislocation density: 1010/cm2 to ~ 107-108/cm2 comparing Figure 15b and Figure 15d, respectively.Figure 16 illustrates the variance of β-phase (bcc) boundary prominence from more rapidly cooling components, and the appearance of the α'-martensitic phase comparing the “window” thicknesses implicit in experimental structures shown in Figure 2.

In Figure 16b the α-phase boundaries are observed to be essentially free of β (bcc), while the contrast conditions show elongated dislocations within the α-phase. Figure 16a shows much less β-phase composing the boundary zones separating the α-phase in contrast to solid monolithic builds implicit in Figures 15a and c.Finally, we should briefly discuss the mechanical behavior (including residual hardness and tensile properties) measured for a wide range of EBM-manufactured components: including variously built solid monoliths and mesh or foam structures. From the contrasting hardness values indicated in the various figure captions, it can be observed that the Vickers microindentation hardness for solid monoliths can vary from 3.6 GPa at the top (Figure 14d) to between 3.9 and 4.6 GPa at the bottom (Figure 14e) (Murr, et al. 2009a).

2009a). Correspondingly, the average ultimate tensile strength (UTS) and corresponding elongations for solid cylinders have been shown to vary from roughly 1.1 GPa to 1.3 GPa and 12 to 25%, respectively (Murr, et al, 2009a). Figure 17 shows fracture surfaces corresponding to triple-melt-pass-built cylinders tested in tension to failure; with a UTS of ~1.1 GPa and an elongation of 12%. Note that the ductile dimple sizes vary from ~25 µm to 200 µm in diameter, in contrast to the bubble sizes shown in Figure 7c (as well as Figure 7 a and b) which vary from ~20 µm to <200 µm. Consequently there is little if any significance for these retained bubbles in initiating fracture.Figure 12. Three Ti-6Al-4V foam prototypes (a) and an SEM view of the center foam prototype (b). The foam prototypes in (a) correspond to (from left) 0.58, 0.68, and 0.83 g/cm3 densities, respectively.Figure 13. (a) SEM views of cut and polished cellular bridges in the center foam (0.68 g/cm3) in Figure 12(a).

Murr, et al. (2009b) observed that the UTS for α'-phase-rich tensile specimens prepared by selective laser melting (SLM) was equal to or slightly greater than the 1.1 GPa to 1.3 GPa noted earlier for EBM built α-phase monoliths, while the corresponding elongations were less than 10%, significantly lower than normal EBM-built solid monoliths. Recent measurements of the stiffnesses (elastic moduli) for mesh and foam structures varied with density, especially the mesh arrays shown in Figure 9, which ranged from a density of 0.78 g/cm3 and a stiffness of 0.92 GPa, to a density of 1.59 g/cm3, and a corresponding stiffness of 6.15 GPa (Murr, et al 2009c).ConclusionsBuild defects, including unmelted or unsintered powder regions can result from small deviations from optimum build parameters such as scan rate and beam current.

No significant difference in build integrity has been observed for single rake versus double/triple rake layer building, while changing from one to three melt scans at optimized build conditions produces increasing α-phase microstructure dimensions and corresponding decreases in residual hardness. It is also observed that in solid, monolithic builds, there is a larger α-phase grain size at the top of the build in contrast to the bottom of the build, and corresponding residual hardness differences (higher hardness at the build bottom versus lower hardness at the top of the build). This results by thermal management differences (or cooling rate variations). Correspondingly, for rapid cooling rates characteristic of small, reticulated mesh arrays and open cellular foams, the microstructures tend toward α'-martensitic platelets which are correspondingly harder than the α- phase acicular platelets as a consequence of their smaller thickness.

Argon bubbles carried into the build from the atomized precursor Ti-6Al-4V powders do not appear to vary with increasing melt scans nor do they noticeably influence the tensile behavior of solid monoliths as observed by fracture surface features dominated by varying sizes of ductile dimples.Acknowledgements This research was supported in part by Mr. and Mrs. MacIntosh Murchison Chair endowments at the University of Texas at El Paso and Lockheed Martin/Aeronautics. Any expressed opinions, findings, conclusions, or recommendations are those of the authors, and do not necessarily represent those of Lockheed Martin/Aeronautics.ReferencesGaytan, S.M., Murr, L.E., Medina, F., Martinez, E., Lopez, M.I., and Wicker, R.B. (2009).Advanced metal powder-based manufacturing of complex components by electron beam melting. Materials Technology, in press.Murr.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceThe Effect of Cell Size and Surface Roughness on the Compressive Properties of ABS Lattice Structures Fabricated by Fused Deposition ModelingL. Mason and M. C. LeuDepartment of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractResearchers looking to improve the surface roughness of acrylonitrile butadiene styrene(ABS) parts fabricated by fused deposition modeling (FDM) have determined that acetone smoothing not only achieves improved surface roughness but increases compressive strength as well. However, the sensitivity of ABS parts to acetone smoothing has not been explored.

However, the sensitivity of ABS parts to acetone smoothing has not been explored. In this study we investigated FDM-fabricated ABS lattice structures of various cell sizes subjected to cold acetone vapor smoothing to determine the combined effect of cell size and acetone smoothing on the compressive properties of the lattice structures. The acetone-smoothed specimens performed better than the as-built specimens in both compression modulus and maximum load, and there was a decrease in those compressive properties with decreasing cell size. The difference between as-built and acetone-smoothed specimens was found to increase with decreasing cell size for the maximum load.IntroductionFused deposition modeling (FDM) is a type of additive manufacturing (AM), or 3Dprinting, where plastic filament is heated and extruded through a nozzle to build up layers of a 3D part.

AM has the capability to quickly create parts with complex geometries that are not possible using traditional methods of manufacturing. This capability has been used in automotive, aerospace, and medical fields to create lightweight designs for many applications. These lightweight designs often incorporate lattice structure elements within them, so it is important to understand as much as possible about lattice structures and how they function in order to use them most efficiently.Lattice structures are engineered to resemble foams with specific properties, but they areknown for generally having an excellent strength-to-weight ratio, and a high surface-area-to- volume ratio. The relative density of a lattice structure, or the ratio of material volume to the total volume the lattice structure occupies, is known as “the single most important structural characteristic” of a lattice structure [1]. As such, relative density has been extensively researched by many academicians.

As such, relative density has been extensively researched by many academicians. In this study, the relative density is kept constant, and another structural characteristic is studied: cell size.To the best of our knowledge, no journal articles have been published on the effects ofvarying cell size of FDM lattice structures while keeping the relative density constant. For open- cell alumina foam, the relative elastic modulus does not change with different cell sizes, but the strut strength decreases with increasing cell size [2]. For body-centered cubic SLM Ti-6Al-4V, the relative elastic modulus and tensile strength decrease with increasing cell size [3]. For gyroid2208SLM 316L stainless steel, the elastic modulus and compression yield strength decrease with increasing cell size [4]. For gyroid DMLS AlSi10Mg, compression strength and microhardness decrease with increasing cell size [5].

For gyroid DMLS AlSi10Mg, compression strength and microhardness decrease with increasing cell size [5]. Many mechanical properties have been maximized by decreasing the unit cell size for metal AM lattice structures, but we do not know if this holds true for FDM fabricated lattice structures.Postprocessing is an important step in many manufacturing processes, but it is especiallyimportant for additive manufacturing. Nearly all AM methods create parts with a high surface roughness, no matter the material. This high surface roughness leads to stress concentrations and is structurally inefficient. Researchers looking to improve the surface roughness of acrylonitrile butadiene styrene (ABS) parts fabricated by FDM have determined that acetone smoothing not only achieves improved surface roughness but increases compressive strength as well [6], [7]. This could be very beneficial to lattice structures, which are known for already having an excellent strength-to-weight ratio.

This could be very beneficial to lattice structures, which are known for already having an excellent strength-to-weight ratio. Improving the compressive strength of ABS lattice structures even further using acetone smoothing could expand applications for plastic lattice structures and improve their uses across the board.Postprocessing of metal AM lattice structures has already gained attention to improve thesurface roughness of the lattice structures. Metal AM structures can be polished using chemical etching when conventional methods such as machining or blasting are not possible, as is the case with lattice structures. Ti-6Al-4V can be etched using an aqueous solution of hydrofluoric acid (HF) and nitric acid (HNO3), which is a subtractive process that takes off an outer layer of material.

Ti-6Al-4V can be etched using an aqueous solution of hydrofluoric acid (HF) and nitric acid (HNO3), which is a subtractive process that takes off an outer layer of material. This postprocessing method typically takes 30 minutes to a few hours to complete and improves the stiffness-to-density ratio [8], [9].Smoothing of ABS parts using acetone is done using highly varying approaches that donot all agree on the same procedures for the “best” outcome. There are two main methods of acetone smoothing: submerging the part in a bath of acetone and subjecting the part to acetone vapors. Acetone bath smoothing is the fastest method, as it involves simply submerging the part in acetone for a few minutes. Submerging a part for 3-7 minutes significantly increases the ductility of the ABS and degrades the tensile strength of the part according to Jayanth et al [10]. Gautam et al.

Submerging a part for 3-7 minutes significantly increases the ductility of the ABS and degrades the tensile strength of the part according to Jayanth et al [10]. Gautam et al. tested kagome lattice structures that had been smoothed using acetone baths, and they found that smoothing increased the compression strength and stiffness, and determined that 5 minutes is the best time for acetone bath smoothing [7].Acetone vapor smoothing can be done using hot or cold vapors. Hot acetone vapor smoothing takes a few minutes, but the acetone is heated to release the vapors quickly. Hot acetone vapor has been known to increase the compressive strength of ABS after smoothing for 5, 7.5, and 10 minutes, although the 5 minutes smoothing withstood the highest compression force [6]. According to Lalehpour et al.

According to Lalehpour et al. [11], three hot vapor baths of 15 s each is the best way to smooth an ABS part to get the best surface roughness.Cold acetone vapor smoothing works through evaporation and therefore takes more time, usually one or more hours, but is much safer than hot acetone vapor smoothing due to acetone’s high flammability. Cold acetone vapor smoothing has not been widely reported in academic literature. One study was found that used cold acetone vapor smoothing. Zhang et al. [12] used2209varying amounts of acetone poured on paper tissues and lined inside a beaker to smooth ABS parts for 30 minutes. The hobby community, however, seems to mainly use cold acetone smoothing to post-process their ABS parts.

The hobby community, however, seems to mainly use cold acetone smoothing to post-process their ABS parts. The cold acetone vapor smoothing method used in our study was based off of an article written by Susi Woods on the website rigid.ink [13].In this study, FDM-fabricated ABS lattice structures of various cell sizes subjected to cold acetone vapor smoothing were investigated to determine the combined effect of cell size and acetone smoothing on the compressive properties of the lattice structures. The acetone- smoothed specimens performed better than the as-built specimens in both compression modulus and maximum load, and there was a decrease in those properties with decreasing cell size.

The acetone- smoothed specimens performed better than the as-built specimens in both compression modulus and maximum load, and there was a decrease in those properties with decreasing cell size. The difference between as-built and acetone-smoothed specimens was found to increase with decreasing cell size for the maximum load.MethodsLattice Structure DesignThe unit cell structure was a macro body-centered cubic (BCC) pattern with strutsconnecting the center to all eight corners of the unit cell cube. The BCC pattern was chosen because it is a commonly used lattice structure that requires no support material to print. The relative density of about 10.5% and specimen size of a 76.2 mm (3 in.) cube were kept constant. The lattice structures were created using the nTopology Element software. A 76.2 mm cube in STL format was imported into the Element software, and the lattice was generated using the cube vertex centroid rule.

A 76.2 mm cube in STL format was imported into the Element software, and the lattice was generated using the cube vertex centroid rule. The struts were then thickened to have a uniform diameter, and a mesh of the lattice structure was generated and exported as an STL file.The parameters of the lattice structures are shown in Table 1. The diameters of the struts were designed to be even multiples of the raster width of 0.508 mm (0.02 in.) so that the layers could be concentric ovals. This was done to minimize gaps and excess material in each layer and to maximize layer stability. The smallest strut diameter possible using this method was 1.016 mm (0.04 in.). The smallest cell size was then chosen to be 6.35 mm (0.25 in.) because that was an easily scaled size and it gave a reasonable relative density of about 10%. The cell size was then doubled, tripled, and quadrupled to obtain the 12.70 mm (0.50 in.), 19.05 mm (0.75 in.), and 25.40 mm (1.00 in.) cell sizes.

The cell size was then doubled, tripled, and quadrupled to obtain the 12.70 mm (0.50 in.), 19.05 mm (0.75 in.), and 25.40 mm (1.00 in.) cell sizes. The smallest overall specimen size that would allow for whole unit cells throughout all the specimens was a 76.2 mm (3 in.) cube.Differences in the relative density, dimensions, and volume of material used are mainlydue to the rounded caps on the outside corners and ends of the lattice structures. These caps ensured that the lattice structures performed appropriately and that the outer corners were not unnecessarily weakened, but they did add more material to the specimens for the larger cell sizes.

The volume and dimensions were taken from the Stratasys Insight software that was used to slice and generate toolpaths for printing the lattice structures.2210Table 1: Lattice Structure ParametersLattice cells3x3x34x4x46x6x612x12x12Cell Size (in.)1.000.750.500.25Cell Size (mm)25.4019.0512.706.35Volume (in.^3)3.4933.2083.1822.904Volume (mm^3)57240525705214447588Dimensions (in.)3.1923.1323.0883.035Dimensions (mm)81.07279.54878.43577.086Strut Diameter (in.)0.160.120.080.04Strut Diameter (mm)4.0643.0482.0321.016Relative Density10.74%10.44%10.80%10.39%ManufacturingFigure 1 shows pictures of one replication. In this study, data from six replications wereanalyzed. There were four cell sizes and two relative roughness values: as-built and acetone- smoothed, for a total of eight treatment combinations.

In this study, data from six replications wereanalyzed. There were four cell sizes and two relative roughness values: as-built and acetone- smoothed, for a total of eight treatment combinations. Each replication of lattice structures included all eight treatment combinations, for a total of 48 specimens. All specimens were printed using the fused deposition modeling (FDM) process on a Stratasys Fortus 400mc using white ABS-M30. One full replication was printed at a time, and the lattice structures were randomly placed in the positions shown in Figure 2 in the 355.6 x 406.4 x 355.6 mm3 (14 x 16 x 14 inch3) build volume to account for possible differences in placement within the build chamber.The rounded feet of the lattice structures did not reliably stick to the support material raftthat is automatically printed under every print in the Fortus.

This made necessary some reinforcing support material under the first struts of every specimen to ensure that the lattice structures printed reliably and well. The support structure used was designed to allow the lattice structures to be broken off the support with minimal harm. The only problem that came of the support structure was the loss of some corner struts on the 6.35 mm specimens. The corner struts are the most fragile, especially on the smallest cell size, since they are only connected to the structure on one end.

The corner struts are the most fragile, especially on the smallest cell size, since they are only connected to the structure on one end. Figure 1d shows only one surviving corner strut out of the four visible lower corners of the 6.35 mm specimens.2211(a)(b)(c)(d)Figure 1: Pictures of one full replication with 8 specimens, with as-built (left) and acetone- smoothed (right) specimens of cell size (a) 25.40mm (1.00 in), (b) 19.05 mm (0.75 in), (c) 12.70 mm (0.50 in), and (d) 6.35 mm (0.25 in)□□□□ □□□□Figure 2: Specimen positions within Fortus 400mc build volume, replication 1Acetone SmoothingHalf of the specimens, one of each cell size in each replication, were subjected to cold acetone vapor smoothing. All four specimens from one replication were smoothed at the same time to facilitate equality within each replication. The specimens were smoothed in a 9 L polypropylene container on top of a polypropylene stage, shown in Figure 3, all within a fume2212hood.

The specimens were smoothed in a 9 L polypropylene container on top of a polypropylene stage, shown in Figure 3, all within a fume2212hood. The stage held one specimen on each corner and had slots machined in between the specimens to allow for improved airflow. A fan underneath the stage circulated the vapors within the container by forcing air down through the stage slots and out and around the specimens. The vapors originated from acetone-soaked paper towels lining the bottom of the container. Each replication was smoothed for 50 minutes using 80 mL of acetone.The smoothing time and amount of acetone to use were determined by testing specimensthat failed to print correctly, such as ones that were interrupted mid-print or ones that did not stick to the printing bed.

First, the amount of acetone was determined by testing in increments of 20 mL and checking every half an hour for two hours; 20 mL did nothing visible for two hours, 40 mL made the specimens sort of glossy, 60 mL started smoothing the specimens but did not finish in two hours, and 80 mL had warped the specimens in an hour and a half. The amount chosen was 80 mL, and then the time to smooth the specimens was narrowed. The specimens were checked every 5 minutes until the smallest lattice structure just started to warp at one hour. Fifty minutes was then chosen as the amount of time to smooth the specimens. The purpose behind this stopping point was to give the acetone as much time as possible to smooth the parts without causing any warping due to loss of structural integrity.When placing the specimens into the container and removing them from the container,only the polypropylene stage was touched.

The specimens were placed on the stage and then the stage was lowered into the container on top of the fan and the lid of the container was closed and sealed. The specimens were removed in a reverse fashion once the smoothing was complete. All specimens were allowed to dry for at least twelve hours in the fume hood before handling. Handling an ABS part directly after acetone smoothing can introduce surface deformations and dust particles that would embed themselves into the malleable surface permanently. At least two weeks passed in between acetone smoothing the specimens and compression testing the specimens to ensure that the ABS had completely resolidified.

At least two weeks passed in between acetone smoothing the specimens and compression testing the specimens to ensure that the ABS had completely resolidified. The mass of each specimen was recorded before and after smoothing.Figure 3: Polypropylene stage for cold acetone smoothing process2213Compression Testing and Surface Roughness MeasurementThe lattice structures were compression tested using an Instron 5969 Universal TestingSystem at a rate of 7.5 mm/min until catastrophic failure. After the specimens were compression tested, an outside edge strut of each specimen was broken off for surface roughness measurement with a profilometer. The profilometer used was a KLA-Tencor P-17, which has a stage with locating holes. To reliably position the broken struts on the stage, a locating fixture, shown in Figure 4, was created and printed from ivory PLA on a Prusa i3 MK2 printer.

To reliably position the broken struts on the stage, a locating fixture, shown in Figure 4, was created and printed from ivory PLA on a Prusa i3 MK2 printer. The two locating pegs on the bottom of the fixture sit in the locating holes on the profilometer stage, while the lattice structure strut sits along the groove on the top of the fixture. Since the strut is cylindrical, the groove ensures that the struts are always aligned with the x-axis of the profilometer. The end of the groove (shown on the left side of Figure 4b) that is in between the two locating pegs serves as a locating point to push the lattice structure strut up against.(a)(b)Figure 4: Locating fixture for profilometer measurements: (a) bottom with locating pegs, (b) top with locating groove for lattice structure strutsOne additional strut from each specimen in the first replication was taken to look at the cross sections of the struts.

The struts were broken off each specimen after compression testing and set in epoxy. In order to stand the struts up in the epoxy mold, the struts were grouped together on top of a bead of epoxy, and another bead of epoxy was dropped onto the cluster of struts and allowed to harden. Then, epoxy was poured around the cluster to complete the mold. The epoxy-set cluster of struts was then polished following the methods described in Metallography by G.

Vander Voort [14] using the following steps, resulting in the cross sections of all eight specimens shown in Figure 5:1) 400 grit SiC paper at 200 rpm with 5 lbs of force for 45 s increments until desired region of specimens was reached2) 600 grit SiC paper at 200 rpm with 3 lbs of force for 45 s 3) 800 grit SiC paper at 200 rpm with 3 lbs of force for 45 s 4) 1200 grit SiC paper at 200 rpm with 3 lbs of force for 45 s 5) 9 μm water-based diamond suspension at 150 rpm with 3 lbs of force for 5 min 6) 3 μm water-based diamond suspension at 120 rpm with 3 lbs of force for 3 min 7) 1 μm water-based diamond suspension at 120 rpm with 3 lbs of force for 3 min2214Figure 5: Cross sections of one strut of each treatment combination from replication 1, with (a) acetone-smoothed 6.35 mm (0.25 inch) cell size, (b) as-built 6.35 mm (0.25 inch) cell size, (c) acetone-smoothed 12.70 mm (0.50 inch) cell size, (d) as-built 12.70 mm (0.50 inch) cell size, (e) acetone-smoothed 19.05 mm (0.75 inch) cell size, (f) as-built 19.05 mm (0.75 inch) cell size, (g) acetone-smoothed 25.40 mm (1.00 inch) cell size, (h) as-built 25.40 mm (1.00 inch) cell sizeResults and DiscussionThe mass of each specimen was recorded before and after smoothing, the averages of which are shown in Table 2.

All the acetone-smoothed specimens increased in mass, but the increase was no more than 5%. There is a definite correlation between the increase in mass and an increase in surface area. The smallest cell size, 6.35 mm, is not the structure with the least mass. This is due to the extra material that was deposited as strands that connected different struts within the same layer and potential further excess material from printing inconsistencies.2215Table 2: Average mass before and after acetone smoothingCell Size (in)Cell Size (mm)As-Built Mass (g)Std. Dev.Acetone- Smoothed Mass (g)Std. Dev.Difference (g)Difference (%)I1.0025.4055.5190.15356.4500.3820.9301.6750.7519.0548.6290.12649.6430.3911.0142.084I0.5012.7045.6770.16947.1920.5071.5153.3150.256.3549.0570.14450.9580.5581.9013.874The cross sections shown in Figure 5 give some insight into how far the cold acetone vapors penetrate the struts.

The struts were built on the diagonal, so there are multiple layers shown in the cross sections. In cross sections e-h, there are two outer rings of material composed by the first and second contours of each layer. A contour is known as the outline of each layer in FDM, counted from the outside of the layer towards the middle, and in this study, there is a maximum of two contours per layer. Only the first contour (the outermost) is affected by the acetone smoothing, and even then, not completely. The ridges on the outer ring of material that can be seen in cross sections f and h are still present on the inside of the outer rings in cross sections e and g. This implies that the acetone vapors were not able to penetrate the full thickness of the outer ring of material.The surface roughness of one strut broken off of each specimen was measured using aprofilometer, and the results are shown in Figure 6.

The profilometer could not record an amplitude of greater than 163.5 microns in the positive or negative direction, which limited the areas of the struts that could be measured. There were sections of each strut that were within these limitations, but the sections that were measured typically had to be the smoothest and most consistent sections on the struts. The Ra values that were obtained are therefore a lower bound estimate instead of an average estimate for the surface roughness of the lattice structures. This being said, the average values that were measured from these specimens do indicate that cold acetone smoothing decreases the surface roughness of ABS specimens fabricated by FDM, as expected. This can also be seen in the cross sections shown in Figure 5.The lattice structures with the two largest cell sizes have the smallest surface roughness values, and of those, the as-built specimens are extremely similar.

The surface roughness of the lattice structures increases for the two smaller cell sizes for both as-built and acetone-smoothed specimens. The range of Ra values for the acetone-smoothed specimens increases significantly as the cell size decreases. The average surface roughness of the smallest cell size is by far the highest at 36.9 microns. This high surface roughness can be attributed to the stability of the printing process. Figure 7 shows an up-close visual comparison between as-built and acetone- smoothed specimens of each cell size, where it is shown that the uniformity of the struts decreases with cell size. This is a symptom of a decreasing number of layers and smaller, less- uniform layers.

This is a symptom of a decreasing number of layers and smaller, less- uniform layers. Smaller layers are not printed as accurately or consistently as larger layers because the printing inconsistencies are amplified for smaller layers and make more of a difference for smaller geometries.2216As-Built■Acetone-Smoothed■50.0) s n o r c i m(40.030.0a R e g a r e v A20.010.025.024.627.436.90.09.58.813.515.025.419.0512.76.35Cell Size (mm)Figure 6: Average surface roughness Ra vs. cell size. The error bars indicate maximum and minimum values(a)(b)(c)(d)Figure 7: Visual comparison between as-built (left) and acetone-smoothed (right) specimens of (a) 25.40 mm (1.00 in), (b) 19.05 mm (0.75 in), (c) 12.70 mm (0.50 in), and (d) 6.35 mm (0.25 in) cell size2217The elastic modulus of each specimen was calculated, and the average elastic modulusfor each treatment combination is shown in Figure 8.

The maximum load was recorded for each specimen, and the average maximum load for each treatment combination is shown in Figure 9. Most of the specimens broke along a perfect 45-degree angle, as expected. The compression results show a definite correlation between decreasing compressive properties and decreasing cell size. This was not expected as Yan et al. [4], [5] claims that the smaller cell sizes of metal lattice structures have better compressive properties than larger cell sizes. For metal AM, smaller cross-sectional areas mean shorter scan distances which leads to faster scanning of adjacent sections and increases the temperature of the smaller scanned area. Increased temperature gives the right conditions for higher compression strength and modulus.Conversely, smaller FDM layers are not printed as accurately or consistently as largerFDM layers.

Increased temperature gives the right conditions for higher compression strength and modulus.Conversely, smaller FDM layers are not printed as accurately or consistently as largerFDM layers. For smaller layers, the machine must move in shorter, faster bursts which introduces printing inconsistencies due to machine backlash and shaking. For larger layers, it takes more time, the movements are smoother, and the direction changes are more spread out. Therefore, the decreasing compressive properties with decreasing cell size in this study can be attributed to the decreasing uniformity and quality of the struts.As-Built■Acetone-Smoothed■) a P M(s u l u d o M6.05.04.03.05.36.04.15.23.8I 5.03.3I -4.2e g a r e v A2.01.00.025.419.0512.76.35Cell Size (mm)Figure 8: Average elastic modulus vs. cell size.

cell size. The error bars indicate maximum and minimum valuesLattice structures are known for their high strength-to-weight ratio, also called specificstrength or strength-to-mass ratio. The strength-to-weight ratio is calculated by dividing the material’s strength by its density. The material’s strength is the maximum load the lattice structure could bear divided by the cross-sectional area of the structure as a whole, which in this case is 5806.44 mm2, or 9 in.2. The material’s density is the mass of the lattice structure in kilograms divided by the volume of the structure as a whole, which in this case is 442451 mm3, or 27 in.3. Specific strength therefore has units of Pa*m3/kg or N*m/kg. The calculated average strength-to-weight ratios are shown in Figure 10.

Specific strength therefore has units of Pa*m3/kg or N*m/kg. The calculated average strength-to-weight ratios are shown in Figure 10. This graph is similar to the maximum load graph in that the difference between as-built and acetone-smoothed specimens increases with decreasing cell size, but it is interesting to note that the smallest cell size had a much lower2218strength-to-weight ratio than the others. This result for the 6.35 mm cell size was a combination of the lowest maximum load and a mass on par with the 19.05 mm cell size. The highest strength-to-weight ratio came from the acetone-smoothed 12.70 mm cell size lattice structure.

The highest strength-to-weight ratio came from the acetone-smoothed 12.70 mm cell size lattice structure. The 12.70 mm cell size lattice structure had a mid-range maximum load and the smallest mass of all the cell sizes, which combined to give it the best acetone-smoothed strength-to-weight ratio.As-Built■Acetone-Smoothed■)N(d a o L x a M2000.01500.01000.01834.1I 1930.81538.51688.61437.4I 1693.11145.51507.0e g a r e v A500.00.025.419.0512.76.35Cell Size (mm)Figure 9: Average maximum load vs. cell size. The error bars indicate maximum and minimum valuesAs-Built■Acetone-Smoothed■o i t a R3000t h g i eW - o t - h t g n e r t S e g a r e v A) g k / m * N(250020001500100050002517260625.42411259219.052734239812.7225317796.35Cell Size (mm)Figure 10: Average strength-to-weight ratio vs.

cell sizeThere was an increase in ductility from the as-built specimens to the acetone-smoothedspecimens. All the as-built specimens made clean breaks when they failed, but some of the acetone-smoothed 25.40 mm (1.00 in.) cell size specimens did not break apart completely when they failed. These specimens fractured and bent a diagonal plane of struts at both joints, but they2219were still held together by the outer layer of material. The outer layer was the material that was most affected by the acetone and exhibited significantly more ductile behavior, while the inner layers were not as affected by the acetone and behaved similarly to the as-built specimens.It can be observed that subjecting the lattice structures to acetone vapor smoothing increases the compressive properties of the lattice structures in general. Both the elastic modulus and maximum load consistently increased from as-built to acetone-smoothed, with little to no overlap.

Both the elastic modulus and maximum load consistently increased from as-built to acetone-smoothed, with little to no overlap. This can be attributed at least in part to the near-elimination of stress concentrations in the acetone-smoothed specimens. Acetone smoothing also increased the effective diameter of the struts by filling in the stress concentrations and indentations where the layers meet. The effective diameter in this case is the smallest diameter of the strut, which happens where two layers meet. The diameter of the contact between the two layers is considered the effective diameter because any material that extends beyond that would not take any of the load held by the strut. After acetone smoothing, the area of contact between two layers is increased by taking the extra material that did not contribute to taking the load before and spreading it evenly along the strut, therefore increasing the effective diameter and using more of the material to carry the load.

This increases the bond between layers by increasing the contact area between them.Table 3 shows the percent difference between the average maximum load results for as-built and acetone-smoothed specimens. It is very apparent that the difference in average maximum load increases as the cell size decreases, meaning that the acetone smoothing had a larger impact on the smaller cell sizes than on the larger cell sizes. This trend is in part caused by the increased uniformity of the struts from acetone smoothing, especially for the smaller cell sizes, and the increased cohesion between layers due to the increased effective diameter.Table 3: Percent difference between average values of maximum loadCell Size (mm)25.4019.0512.706.35Avg. Max Load5.1%9.3%16.3% 27.3%An analysis of variance (ANOVA) was run on each set of compression results to determine if the response from the two combined factors, cell size and surface roughness, was additive or interactive.

For either model, factor a creates a certain independent response, A, and factor b creates another independent response, B. The additive model’s response is just A+B, meaning adding the factors’ separate responses together results in the combined response. The interactive model adds another variable to the combined response equation, called the interaction effect variable, C, making the response equation A+B+C. This interaction effect variable represents the response from the non-additive relationship between factors a and b. A two-factor ANOVA simply determines if the model is interactive or additive, and it does not determine what the interactive effect is. The ANOVA results shown in Table 4 indicate that the interaction between the cell size and qualitative surface roughness is significant for the elastic modulus and the maximum load because the p-value is less than 0.05, meaning the model is interactive for both of those responses.

This interactive model can especially be seen in Figure 9, where the increase in maximum load from as-built to acetone-smoothed changes based on the cell size. If2220this was an additive model, acetone smoothing the specimens would have increased the maximum load by the same amount no matter the cell size.Table 4: ANOVA resultsDataInteraction P-valueModulus0.0002Max Load< 0.0001ConclusionsLattice structures of four different cell sizes, half of which were exposed to cold acetone vapor smoothing, were tested in compression to determine the combined effect of cell size and acetone smoothing. The difference between as-built and acetone-smoothed specimens was found to increase with decreasing cell size for the maximum load. This trend was caused by the increased uniformity of the struts from acetone smoothing, especially for the smaller cell sizes, and the increased cohesion between layers due to the acetone fusing the layers together on the outside.

The acetone-smoothed specimens performed better than the as-built specimens in both elastic modulus and maximum load, and there was a decrease in those compressive properties with decreasing cell size. The increase in compressive properties for the acetone-smoothed specimens can be attributed at least in part to the reduction of stress concentrations and the increase in effective diameter. The decreasing compressive properties with decreasing cell size can be attributed to the decreasing uniformity and quality of the struts. There was also an increase in ductility from the as-built specimens to the acetone-smoothed specimens in the outer layer of material. Overall, the acetone smoothing affected the compressive properties of the smaller cell sizes more significantly than the larger cell sizes. It was determined through an ANOVA test that the two factors, cell size and qualitative surface roughness, were not purely additive and that they interacted to give unique results.

It was determined through an ANOVA test that the two factors, cell size and qualitative surface roughness, were not purely additive and that they interacted to give unique results. Therefore, there exists an optimum combination of cell size and surface roughness that gives the best response depending on the application.References[1]L. J. Gibson and M. F. Ashby, Cellular Solids: Structure and Properties, Second Edi. Cambridge University Press, 1999.[2] R. Brezny and D. J. Green, “The effect of cell size on the mechanical behavior of cellularmaterials,” Acta Metall. Mater., vol. 38, no. 12, pp. 2517–2526, 1990.[3]I. Maskery, A.O. Aremu, M. Simonelli, C. Tuck, R.D. Wildman, I.A. Ashcroft, and R.J.M. Hague , “Mechanical Properties of Ti-6Al-4V Selectively Laser Melted Parts with Body-Centred-Cubic Lattices of Varying cell size,” Exp. Mech., vol. 55, no. 7, pp. 1261– 1272, 2015.[4] C. Yan, L. Hao, A. Hussein, and D.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of Constrained Surface Texturing on Separation Force in Projection StereolithographyHaiyang He1, Yayue Pan *1, Jie Xu1, Xiaoming Yu21 Department of Mechanical and Industrial Engineering, University of Illinois at Chicago, Chicago, IL, 60607, USA2 College of Optics &Photonics, University of Central Florida, Orlando, FL, 32816, USAAbstractIn projection stereolithography (SL) processes, the separation of a newly cured layer from the constrained surface is a historical technical barrier and still greatly limits its printable size, process reliability and print speed. This paper presents an approach to reduce the separation force in projection stereolithography (SL) processes by texturing the constrained surface with radial microgroove patterns.

This paper presents an approach to reduce the separation force in projection stereolithography (SL) processes by texturing the constrained surface with radial microgroove patterns. Separation forces with conventional smooth constrained surface and textured surface are both modeled. The analytical model suggests that a proper design of micro patterns of the constrained surface is capable of reducing separation forces greatly. Furthermore, a projection SL testbed with online separation force monitoring unit is developed for experimental study. Experimental results verified the effectiveness of micro surface textures in reducing separation forces. Test cases also show that with the help of the proposed textured constrained surface, parts with wide solid cross sections that could not be printed using conventional methods were manufactured successfully.

The influence of the textured constrained surface on the printed parts’ surface roughness is studied, a grey scale projection approach is proposed to eliminate the influence and improve the surface quality of printed parts. Hence, the presented methods can help to improve the manufacturing capability of Projection SL processes.Keywords: Additive Manufacturing, Projection Stereolithography, Surface Texture, Constrained Surface.IntroductionProjection Stereolithography (SL) is one of the most important Additive Manufacturing technologies currently available and also the first commercialized AM technology. In Projection SL process, liquid photosensitive polymer is cured usually through the use of a DLP projector, which supplies the amount of energy via projecting digital mask images to induce a curing reaction, forming a highly cross-linked polymer.

Compared to other polymer additive manufacturing techniques like extrusion or jetting processes, Projection SL produces parts with the highest accuracy and the best surface finish.There are two ways to cure liquid polymer in Projection SL, free surface method and constrained surface method, as illustrated in Fig.1. Compared to free surface process, the constrained surface process has several advantages including less material waste, higher resolution and faster speed, and is being used increasingly recently.

Compared to free surface process, the constrained surface process has several advantages including less material waste, higher resolution and faster speed, and is being used increasingly recently. Despite the advantages, the separation of a newly cured layer from the constrained surface is a historical technical barrier and still greatly limits the printable size, process reliability and print speed [1].11735ComputerResinCured partConstrained surfaceTankPlatformResinTankPlatformResinTankPlatformTextured surfaceComputer Digital Micromirror DeviceCured partTextured surfaceTextured surfaceComputer Digital Micromirror DeviceCured partDigital Micromirror DeviceFigure 1. A schematic diagram of constrained surface based Projection SLAs illustrated in Fig. 1, in constrained surface Projection SL process, the separation force occurs in the process of lifting newly cured layer up and refilling liquid resin into the gap created during lifting process.

1, in constrained surface Projection SL process, the separation force occurs in the process of lifting newly cured layer up and refilling liquid resin into the gap created during lifting process. Hence, the separation force exerted on newly cured layer is actually the suction force caused by pressure difference in liquid filling[2].Extensive research has been conducted to reduce separation forces. Approaches like peeling, two-way movement, and coating have been developed. However, these approaches may prolong total the build time, require much larger liquid tank, and frequent replacements of coatings [3, 4]. The peeling approach introduces other problems like corrugated constrained surfaces and large peeling force [5, 6]. These drawbacks are more significant when the print area is large. Hence separation problem still remains to be the primary challenge and limits the manufacturing capability of Projection SL.

These drawbacks are more significant when the print area is large. Hence separation problem still remains to be the primary challenge and limits the manufacturing capability of Projection SL. For example, it is worth noting that building envelopes of current Projection SL machines are all in inch level, and no greater than 8 L x 5 W x 10 H. To produce a part in feet level is almost impossible. Therefore, it is of highly significance to reduce the separation force and enhance the capability of Projection SL.Against this background, this paper aims to develop novel approaches to reducing separation forces, without sacrificing other properties like speed, volume and reliability. A constrained surface texturing approach for separation force reduction is developed in this study.1.

A constrained surface texturing approach for separation force reduction is developed in this study.1. Effects of Constrained Surface Texture on Separation ForceIn constrained surface SL, conditions of the interfaces, the newly cured layer surface and constrained surface, play important roles in determining separation forces. In addition, surface texturing has been widely used for altering pressure gradient or liquid flow dynamics in various applications [7, 8, 9, 10, 11], but not yet in constrained surface design for stereolithography. Hence, in this study, we investigate how to modify the constrained surface to reduce separation forces in SL process.17362Top viewRΔrrNewly cured layerhVelocity profileUnit circleNon-textured surfacehNewly cured layerTextured surfaceθwαTop viewVelocity profiled(a)(b)Figure 2.

Schematic of separation areas with a solid circular layer and (a) non-textured constrained surface (b) textured constrained surfaceFig. 2 (a) shows a schematic diagram of the separation area in conventional Projection SL systems based on smooth constrained surface. R is the radius of the part, and h is the height of the gap between the newly cured layer and the coating. Fig. 2(b) is a schematic of separation area with proposed constrained surface, which is patterned with a radial microgroove pattern.

Fig. 2(b) is a schematic of separation area with proposed constrained surface, which is patterned with a radial microgroove pattern. The microgroove pattern has four design parameters: groove width w, groove depth d, the angle of each groove is α and groove radial distance 𝜃.To investigate effects of surface texture on separation forces and learn how to design parameters, the liquid filling process and separation force are modelled for both conventional smooth surface and textured surface.Separation force with smooth constrained surface:By applying Navier–Stokes equations, we can get:𝝁𝝏𝟐𝒖 𝝏𝒛𝟐 =𝒅𝑷𝒅𝒓(1)where 𝜇 is the viscosity of the liquid.

Based on Mass Conservation Equation, the relationship between the Z stage moving speed V and liquid flow rate Q can be found:𝒉 𝟎𝑸 = 𝑽 ∙ 𝝅 ∙ 𝒓𝟐 = 𝟐 ∙ 𝝅 ∙ 𝒓 ∫ 𝒖𝒅𝒛(2)By substituting (1) into(2), the pressure gradient ∆𝑝 and resulted pressure P could bederived as:𝑷 = −∆𝒑 = 𝑽µ𝒓 ∙𝟑µ𝑽𝒉𝟑 ∙ 𝒓𝟐 +𝟔𝒉𝟑 (3) 𝒉𝟑 ∙ 𝑹𝟐 (4)𝟑µ𝑽Thus, the separation force F exerted on part can be calculated by integrating the pressureacross the area:17373𝑹 𝟎𝑭 = ∫ 𝟐𝝅𝒓𝑷𝒅𝒓 =𝟑𝝅∙µ𝑽𝟐∙𝒉𝟑 ∙ 𝑹𝟒 (5)Separation force with textured constrained surface:According to [12], the hydraulic resistance 𝑹𝒉𝒚𝒅 for straight channels with arbitrary cross-sectional shape could be approximated by:𝑷𝒆𝟐 𝑨𝟑 (6) where L is the channel length, 𝝁 the dynamic viscosity of the liquid, A the cross-sectional area and Pe the perimeter.

Also, by definition, the hydraulic resistance is:𝑹𝒉𝒚𝒅 ≈ 𝟐 ∙ 𝝁 ∙ 𝑳 ∙𝑹𝒉𝒚𝒅(𝜟𝒑) ≡𝜟𝒑 𝑸(𝜟𝒑)(7)where 𝜟𝒑 is a certain pressure drop and 𝑸 is the flow rate.From (6) and (7), we can obtain:∆𝒑 =𝒅𝒑 𝒅𝒓= 𝑸 ∙ 𝟐𝝁 ∙𝑷𝒆𝟐 𝑨𝟑 (𝟖)For situation with micro textures,𝒘 𝟐 𝑨 = 𝟐𝝅 ∙ 𝒓 ∙ 𝒉 + 𝟎. 𝟓𝒏 ∙ 𝒘 ∙ 𝒅 (10) 𝑸 = 𝑽 ∙ 𝝅 ∙ 𝒓𝟐 (11) where 𝑽 is the separation speed, 𝒏 the number of grooves. 𝒘 and 𝒅 are the width and depth of the grooves. 𝒉 denotes the height of the gap.𝒅 𝐜𝐨𝐬(𝜶/𝟐)𝑷𝒆 = 𝟐 ∙ 𝟐𝝅 ∙ 𝒓 + 𝟐 ∙ 𝒏 ∙ (−) (9)By substituting (9), (10) and (11) into(8), we can obtain𝒅𝒑𝒅𝒓= 𝑽 ∙ 𝝅𝒓𝟐 ∙ 𝟐𝝁 ∙𝒘 𝟐 (𝟐𝝅∙𝒓∙𝒉+𝟎.𝟓𝒏∙𝒘∙𝒅)𝟑𝒅 𝒄𝒐𝒔(𝜶/𝟐)(𝟐∙𝟐𝝅∙𝒓+𝟐∙𝒏∙ (−𝟐 ) )(12)By taking integral of the pressure 𝑝 on each unit circle, the separation force could be calculated. To investigate the effects of microgrooves on separation force reduction, pressure gradient is calculated for textured surface using Equation (12), and then compared with pressure gradient for smooth surface in Equation (3).

It was found that textured surface will reduce pressure drop by ~60% compared with conventional smooth surface when other conditions are restricted the same and hence smaller separation force than those of smooth PDMS.To achieve smallest separation force, the pressure drop should be lowest. Thus, by takingthe partial derivative of the pressure drop with respect to w in equation (12), we can get:𝝏∆𝒑 𝝏𝒘= 𝟐𝛍𝑽𝝅𝒓𝟐[−𝟐𝐧 (𝟒𝝅𝒓 +𝟐𝒏𝒅 𝐜𝐨𝐬 (𝜶/𝟐)− 𝒏𝒘) (𝟐𝝅𝒓𝐡 + 𝟎. 𝟓𝒏𝒅𝐰) − 𝟏. 𝟓𝒏𝒅 (𝟒𝝅𝒓 +(𝟐𝝅𝒓𝒉 + 𝒏𝒘𝒅)𝟒𝟐𝒏𝒅 𝐜𝐨𝐬 (𝜶/𝟐)𝟐 − 𝒏𝒘)](13)41738Since w, d, n and h are all larger than zero and 𝟒𝝅𝒓 >> 𝐧𝐰, above equation cannot bezero.Similarly, take the partial derivative of the pressure drop with respect to d in equation (12), 𝝏∆𝒑 𝝏𝒅= 𝟐𝝁𝑽𝝅𝒓𝟐𝟒𝒏 𝒄𝒐𝒔 (𝜶/𝟐)(𝟒𝝅𝒓 +𝟐𝒏𝒅 𝒄𝒐𝒔 (𝜶/𝟐)− 𝒏𝒘) (𝟐𝝅𝒓𝒉 + 𝟎. 𝟓𝒏𝒘𝒅) − 𝟏.

𝟓𝒏𝒘𝒅) − 𝟏. 𝟓𝒏𝒘 (𝟒𝝅𝒓 +(𝟐𝝅𝒓𝒉 + 𝒏𝒘𝒅)𝟒𝟐𝒏𝒅 𝒄𝒐𝒔 (𝜶/𝟐)− 𝒏𝒘)𝟐(14)Set𝝏∆𝒑𝝏𝒅= 0𝒅 =𝟏𝟔𝝅𝒓𝒉 + 𝟑𝒏𝒘𝟐𝒄𝒐𝒔 (𝜶/𝟐) − 𝟏𝟐𝝅𝒓𝒘𝒄𝒐𝒔(𝜶/𝟐) 𝟐𝒏𝒘(15)Therefore, the micro texture height and width, d and w should be designed based on therelationship described in equation (15) for achieving the smallest separation force.Based on the study in [13], the oxygen inhibition layer thickness above a PDMS film is ~2 µm. Thus, when h is fixed at 2 µm, the relationship between the width and depth of the micro textures is shown in Figure 3. The relationship above can be used as a guide when designing the parameters of the textures on PDMS.100806012020 Manufacturing a 20 mm diameter cylinder4020Figure 3. The relationship between width and depth of micro textures for achieving the smallest separation force2. Textured Constrained Surface based Projection SL Experimental Setup17395As shown in Fig. 4, a constrained surface based Projection SL testbed is built.

Textured Constrained Surface based Projection SL Experimental Setup17395As shown in Fig. 4, a constrained surface based Projection SL testbed is built. The setup consists of an imaging unit, a resin vat which is an optic clear petri dish, a linear actuator that elevates the build platform, a control board, and a load cell that measures the separation force. The load cell is mounted directly on the platform, so real-time force can be measured and recorded. An off-the-shelf projector is used as the imagining unit. The optical lenses of the projector were modified to reduce the projection distance. Various projection settings, including focus, key stone rectification, brightness, and contrast were adjusted to achieve a sharp projection image on the projection plane. A process control testbed has been developed using C++ language. It integrates the geometry slicing, image projection, and motion controlling.

A process control testbed has been developed using C++ language. It integrates the geometry slicing, image projection, and motion controlling. In addition, an online force monitoring testbed has been developed in Matlab. It reads and processes data from the load cell, and saves the separation force in real time.Z stageResinvatPower supplyplatformsoftwareControl boardLoad cellFigure 4. Hardware SetupConventional Projection SL systems are usually based on smooth Polydimethylsiloxane (PDMS) films or smooth TEFLON films. In this study, PDMS films are prepared and coated on bottom surfaces of the resin vat. To explore the potential of textured constrained surface in reducing separation forces, both conventional smooth PDMS films, and PDMS films with designed surface textures are tested.3. Experimental Results and Discussions3.1. PDMS Texturing and Force MeasurementsFig.

Experimental Results and Discussions3.1. PDMS Texturing and Force MeasurementsFig. 2 (b) is an illustration of the resin replenishment with a textured surface, where the constrained surface is patterned with radially symmetric microgrooves. Those microgrooves will help increase the flow pathway cross section, hence the resin flow rate. Micro patterns with17406different parameter settings, i.e. width and depth of the microgrooves, radius of the centre hole, have been prepared on 1mm thick PDMS substrates. Microscopic images of the top view of four samples are shown in Fig. 5, where the red scale bar is 0.2mm. Surface #1 was textured by micro- milling process, consisting of 8 radially symmetrical microgrooves, with 80µm width and 80µm depth. The other three surfaces were textured by using femtosecond (fs) laser micromachining technology, which utilized a 40-fs laser beam to fabricate the designed microgrooves with high precision.

The other three surfaces were textured by using femtosecond (fs) laser micromachining technology, which utilized a 40-fs laser beam to fabricate the designed microgrooves with high precision. Details of the laser micromachining process can be found in [14]. The geometries of the microgrooves in these three samples are: #2-width 43 µm, depth 46 µm, angle 15⁰; #3-width 60 µm, depth 107 µm, angle 15⁰; #4-width 66 µm, depth 290 µm, angle 15⁰.#4#1234#2#3#4#3#1234#2Figure 5. Microscopic images of four textured surfaces. Scale bar: 0.2mm.Experiments were carried out to test the effectiveness of micro textures on separation force reduction. A solid cylinder with a radius of 10 mm has been built with the conventional Projection SL based on smooth constrained surface, and also with the proposed Projection SL based on the textured surface (#2-width 43 µm, depth 46 µm) as shown in Fig. 5. In all experiments, a velocity of Z elevator of 1.56 mm/s.

5. In all experiments, a velocity of Z elevator of 1.56 mm/s. an acceleration of Z elevator of 1.25 mm^2/s, and a layer thickness of 203.2 um are used. In addition, the same resin MakerJuice G+ with a viscosity of 90 cP is used in all tests.Forces during printing in the developed SL testbed, with the two types of constrained surfaces, conventional smooth PDMS and textured PDMS surfaces, are measured. Fig. 6 plots forces in the process of printing 8 successive layers by using smooth surface and the four textured surface samples in Fig. 5. Compared to the conventional smooth PDMS surface, separation forces in the printing process are reduced by ~60% by using textured surfaces.08060402024 Force (N)……200.5 1.5Figure 6. Separation forces during printing 8 layers by using non-textured PDMS film and the four textured PDMS films showed in Fig. 5.17417Note that the PDMS surface also deforms in the separation forces.

5.17417Note that the PDMS surface also deforms in the separation forces. Hence the peak of separation force doesn't occur exactly in the initial moment of part elevation, and the actual separation velocity V is smaller than the Z stage velocity due to the deformation of PDMS surface. The experimental results verified that the separation force could be reduced greatly by modifying the constrained surface with radial microgroove textures.3.2 Test Case 1: Fabrication of A Hand Model with Textured Constrained Surface Projection SL TechnologyThe effectiveness of textured surface on reducing separation forces were verified by comparing measured separation forces. To further validate the capability of Projection SL based on textured surfaces, multiple complicated structures have been fabricated using the developed testbed and textured surface sample #2. A hand model is printed using proposed approach.

A hand model is printed using proposed approach. Its CAD model, fabricated picture, cross section contour and the corresponding microscopic image are shown in Fig. 7. The scale bar is 5 mm.(a) CAD model (b) Fabricated part(c) CAD contour of the wrist(d) Microscopic image of the fabricated wrist contour5mm(a) CAD model (b) Fabricated part(c) CAD contour of the wrist(d) Microscopic image of the fabricated wrist contour5mmFigure 7. A hand model printed with textured constrained surface.3.3 Test Case 2: Fabrication of A ring model with Textured Constrained Surface Projection SL TechnologyTo validate the capability of the proposed approach for manufacturing parts with delicate features, a ring model is printed using textured sample #2. Its CAD model, fabricated picture, cross section contour and the corresponding microscopic image are shown in Fig. 8.

Its CAD model, fabricated picture, cross section contour and the corresponding microscopic image are shown in Fig. 8. The scale bar is 5 mm.17428(c) CAD contour segment of the ring (d) Microscopic image of fabricated part contour5mm(a) CAD model (b) Fabricated partFigure 8. A ring model printed with textured constrained surfaceAlthough the conventional smooth PDMS surface can also print such parts of this size, the separation force will be much larger than that of textured PDMS, which will result in shorter life of the constrained surface. As the building size increases, there is a high risk that the increasing separation force will damage the parts or even fail the build process.3.4 Test Case 3: Fabrication of a 75 mm diameter cylinder using Textured Constrained SurfacePrinting geometry with a wide solid cross section has always been a challenge for constrained surface SL due to over-large separation force.

To validate the effectiveness of textured surface PDMS constrained surface on fabricating objects with wide solid cross sections with small separation force, a 75mm diameter cylinder, was tested. The height of the cylinder is 25mm.The material used for experiments is MakerJuice G+ (green). The layer thickness is set to be 50 um. The part was printed both in a commercial machine with smooth constrained surface and our prototype with textured constrained surface. With the commercial machine, the part failed during the printing process at around 1 mm in height. The failed part is shown in Fig. 9 (a). However, with the textured constrained surface, the printing job was successfully completed and the printed part is shown in Fig. 9 (b). The comparison between the failure part and our successful part is given in Fig. 9 (c).1743925.4 mm25.4 mm25.4 mm25.4 mm(a)(b)(c)Figure 9. (a) Failure of printing a 75mm diameter cylinder using commercial machine.

9 (c).1743925.4 mm25.4 mm25.4 mm25.4 mm(a)(b)(c)Figure 9. (a) Failure of printing a 75mm diameter cylinder using commercial machine. Part was pulled down from the platform due to large separation force and stick on the constrained surface (b) Success of Printing a 75mm cylinder using a textured constrained surface based testbed. (c) Comparison between the failed part stick on the constrained surface of the commercial machine and successfully printed part using the textured constrained surfaceThe separation force for building the part with textured surface was recorded and is plotted in Fig.10. The average separation force is ~18N for each layer. Besides, there is no significant increase of separation force over the manufacturing process, which is often seen in conventional constrained surface SL processes [15].20003000250050010001500010 Separation Force for Printing a 75mm diameter Cylinder3500Figure 10.

Separation force for printing a 75mm diameter cylinder using the textured constrained surface (~520 layers)The cylinder’s CAD model, fabricated picture, cross section contour and the correspondingmicroscopic image are shown in Fig. 11.(c) Microscopic image(a) CAD model (b) Fabricated part25.4 mm1mm1mm25.4 mm(a) CAD model (b) Fabricated part(c) Microscopic imageFigure 11. A 75 mm diameter cylinder printed with textured constrained surface1744103.5 Test Case 4: Fabrication of Gear using Textured Constrained SurfaceTo further demonstrate the capability and versatility of textured constrained surface, a common mechanical gear (~68 mm diameter, 10mm in height) is printed. The material used is SL600M, produced by Envision TEC. The part was printed using both a commercial machine with conventional constrained surface and our setup with textured surface.

The material used is SL600M, produced by Envision TEC. The part was printed using both a commercial machine with conventional constrained surface and our setup with textured surface. After printing a few layers in the commercial machine, the part was pulled down from the platform by the large separation force and adhered to the constrained surface. Debris could be observed on the metal platform and the constrained surface, as shown in Fig. 12 (b).25.4 mm25.4 mm25.4 mm(a)(b)(c)Figure 12. (a) A CAD model of a mechanical gear (b) Failure of printing a mechanical gear in a commercial machine which uses the conventional non-textured constrained surface (c) Comparison between the failed part printed in the commercial machine and the successfully printed part in our setup using textured constrained surfaceHowever, with the textured surface, the gear can be successfully printed. The comparison between these two parts are given in Fig. 12 (c) (scale bar: 25.4mm).

The comparison between these two parts are given in Fig. 12 (c) (scale bar: 25.4mm). The separation force for manufacturing this gear with textured was recorded during the whole process, as plotted in Fig. 13.20 Separation Force for Printing a Gear60010001200080040040200Figure 13. Separation force for printing the mechanical gear using the textured constrained surface (~200 layers)174511The gear’s CAD model, fabricated picture, cross section contour and the correspondingmicroscopic image are shown in Fig. 14.(a) CAD model (b) Fabricated part(c) Microscopic image1mm25.4 mm(c) Microscopic image(a) CAD model (b) Fabricated part25.4 mm1mmFigure 14. A gear printed with textured constrained surface4.

A gear printed with textured constrained surface4. Discussion of Texture Imprinted on Part SurfaceAlthough the textured constrained surface significantly reduced the separation force and enables successful prints of objects with wide solid cross sections, the textured constrained surface brings a surface finish problem. The textures were imprinted on the part surface and can be observed. To better understand the effects of textured constrained surface on surface finish of printed parts, this section characterizes the part surface finish and investigate approaches to eliminate the imprinted textures.Eight samples were manufactured using the textured PDMS constrained surface, and two different resins, MakerJuice and SL500. Figure 15 (a) shows a sample and Fig. 15 (b) shows a microscopic image of the region that was printed by the micro-groove region of the PDMS film. As shown in Fig.

(b) Scan data of region in Fig. (d) (e)406001400Figure 15. Samples printed with and without gray scale images: (a) a sample printed with the conventional mask image projection method; (b) microscopic image of the imprinted texture on the sample in (a); (c) a sample printed with the proposed gray scale image projection method; (b) microscopic image of the region where microtexture was imprinted; (e) Z profiles of the region of the part surface where was in contact with the micro-groove on the textured constrained surface.To address this texture printing problem, a gray scale projection method was investigated. After curing a layer, a grey scale image (R: 110, G: 110, B: 110) of that layer is projected. Due to the oxygen inhibition adjacent to the constrained surface, only a thin film of resin will be cured on the layer in the non-textured regions. To test the feasibility of this method, the same square sample was printed using this gray scale image method.

To test the feasibility of this method, the same square sample was printed using this gray scale image method. Picture and microscopic image of the sample printed using the gray scale image method are shown in Fig. 15 (c) and (d). The profile of the line region same as the region in Fig. 15 (b) which was contacted with the microgroove of the constrained surface, is plotted as the red curve in Fig. 16. It can be seen from Fig. 15 (d) and (e) that with the gray scale image method, the imprinted texture feature could be eliminated. It indicates that effects of the microgroove in the PDMS constrained surface on the surface finish of the printed part could be eliminated by the gray scale image method. Additionally, the surface roughness of 75 mm diameter cylinder, mechanical gear and square samples fabricated with textured PDMS surface were measured, as listed in Table 1.

Additionally, the surface roughness of 75 mm diameter cylinder, mechanical gear and square samples fabricated with textured PDMS surface were measured, as listed in Table 1. It can be seen that the gray scale image method is a feasible approach to eliminating the imprinted texture and improving the surface roughness of printed parts.174713Table 1. Roughness measurements of parts fabricated with textured constrained surfaceModels Ra (um) Rq(um) Rz(um) Square sample (before eliminating textures) 0.58167 0.39486 23.754 0.25126 0.33688 2.224 Square sample (after eliminating textures) 0.792 75 mm diameter cylinder 0.98444 24.055 0.27684 0.35599 25.155 Mechanical gear5. ConclusionsA novel constrained surface texturing method for separation force reduction in projection stereolithography systems is developed in this study.

ConclusionsA novel constrained surface texturing method for separation force reduction in projection stereolithography systems is developed in this study. Analytical models indicate that by modifying the constrained surface with radial microgroove patterns, the separation force during printing process could be reduced greatly. Several PDMS samples with microgroove patterns are prepared using micro milling and femtosecond laser micromachining technologies. Separation forces during the printing process using textured constrained surfaces are measured and compared with the forces during printing process using the conventional non-textured constrained surface. Multiple parts have been built successfully with the developed Projection SL process based on textured PDMS constrained surfaces. Experimental results verified the effectiveness of constrained surface texturing approach in reducing separation forces and fabricating 3D objects.

Experimental results verified the effectiveness of constrained surface texturing approach in reducing separation forces and fabricating 3D objects. Furthermore, this study demonstrated advantages of the proposed textured constrained surface in Projection SL on fabricating parts with large solid cross section areas, which are very challenging or even impossible with current constrained surface Projection SL technology due to over-large separation forces.Future work will be conducted to further optimize surface texture design, and develop the gray scale image method for eliminating the texture imprinting problem for printing complicated geometries with high surface finish.AcknowledgementThis material is based upon work supported by the National Science Foundation under Grant No. 1563477.

1563477. The authors would like to also acknowledge the Nanotechnology Core Facility of the University Of Illinois at Chicago campus for training on and usage of the Bruker-Nano Model Contour GT-K Optical Profilometer.References[1] K. Ikuta et al. "Real three dimensional micro fabrication using stereo lithography and metal molding." MEMS: 1993, 42-47 [2] H, He et. al. Effect of Surface Texturing on Separation Force in Projection Stereolithography, 11th International Conference on Micro Manufacturing, ICOMM 2016, Irvine, California, March 29-31, 2016 [3] Denken Corp., SLP-4000 Solid Laser Diode Plotter, Product Brochure, Denken Corporation, Japan, 1997. [4] Y. Pan et. al, “A Fast Mask Projection Stereolithography Process for Fabricating Digital174814Models in Minutes.” J. Mfg. Sci. Eng., Vol. 134, No. 5, pp. 051011, 2012 [5] P.M. Lambert et al., "Design Considerations for Mask Projection Microstereolithography Systems." SFFS: 2013, 111-130. [6] J.

Mfg. Sci. Eng., Vol. 134, No. 5, pp. 051011, 2012 [5] P.M. Lambert et al., "Design Considerations for Mask Projection Microstereolithography Systems." SFFS: 2013, 111-130. [6] J. Hendrik et al., "Apparatus and method for the non-destructive separation of hardened material layers from a flat construction plane." U.S. Patent 7,438,846, issued October 21, 2008. [7] D. J. Kukulka et al. "Effect of rigidized textured surfaces on heat transfer and fouling of enhanced tubes," Chem. Eng. Trans., 2009; 18: 575-580. [8] S. Srinivasan et al."Sustainable Drag Reduction in Turbulent Taylor-Couette Flows by Depositing Sprayable Superhydrophobic Surfaces," Phys. Rev. Lett., 2015; 114: 014501. [9] L. Xu, "Liquid drop splashing on smooth, rough, and textured surfaces," Phys. Rev. E, 2007; 75(5): 056316. [10] D. Quere et al. "Slippy and sticky microtextured solids," Nanotechnology, 2003; 14(10: 1109-1112. [11] E. M. Nicolaiewsky et al. "Liquid flow over textured surfaces. 1. Contact angles," Ind. Eng.

FECT OF DEFECTS ON FATIGUE TESTS OF AS-BUILT TI-6AL-4V PARTS FABRICATED BY SELECTIVE LASER MELTINGHaijun Gong*, Khalid Rafi*, Thomas Starr†, Brent Stucker*Department of Industrial Engineering, †Department of Chemical Engineering, J. B. Speed School of Engineering, University of Louisville, Louisville, KY 40292AbstractDefects can be found in parts made using Selective Laser Melting (SLM) due to balling effects and other types of localized irregularities. This study investigates how defects affect the fatigue performance of Ti-6Al-4V samples in an SLM as-built surface finish condition. Fatigue samples were built and heat treated for stress relief. In order to investigate the effect of defects, a series of fatigue samples were designed with built-in cylindrical and double-conical defects. Tests were carried out to correlate maximum stress to the number of cycles to failure.

Tests were carried out to correlate maximum stress to the number of cycles to failure. Optical and scanning electron micrographs were utilized to compare and analyze crack initiation and propagation characteristics. Based on the results, the influence of defects on fatigue properties is discussed.IntroductionAdditive Manufacturing (AM) refers to the process of joining materials to make objects from three-dimensional Computer Aided Design data, usually layer upon layer, as opposed to subtractive manufacturing methodologies, such as traditional machining [1]. It significantly simplifies the process of producing complex 3D objects directly from CAD data without process planning. The operator only needs some basic dimensional details and a small amount of understanding as to how the AM machine works and the materials that are used [2].

The operator only needs some basic dimensional details and a small amount of understanding as to how the AM machine works and the materials that are used [2]. Application of additive manufacturing technology to fabricate complex three-dimensional components is one promising direction in industrial and medical field [3, 4]. This approach is rapidly changing the pattern of designers and processers for creating objects with desired shape and structure. Many additive manufacturing methods have been applied in many variations to satisfy industrial and medical needs for rapidity and flexibility.Titanium alloys such as Ti-6Al-V4 are widely adopted in aerospace, biomedical and industrial fields due to its inherent properties of fracture resistance, fatigue behavior, corrosion resistance and biocompatibility [5].

In recent years, selective laser melting (a metal powder bed fusion AM process which utilizes a laser to melt metal powder layer-by-layer.) has shown great versatility for fabricating parts from numerous types of metallic powders such as stainless steel, maraging steel, cobalt chromium and Ti-6Al-4V. Titanium alloys are extensively investigated and utilized in SLM processes. Thus, there is interest in quality attributes of SLM Ti-6Al-V4 components as compared to wrought or cast plus post heat treatment-produced components [6-8]. In order to give more insights to designers who are employing this technology, quality and life span of AM components must be carefully investigated, especially for metallic parts with defects. Thus, this research focuses on understanding how defects affect the fatigue life of as-built Ti-6Al-4V specimens fabricated using SLM. The stress-life plot is discussed based on the crack pattern of499the fracture surfaces.

The stress-life plot is discussed based on the crack pattern of499the fracture surfaces. Fractography of each crack pattern is shown for interpreting the crack initiation and propagation process.Material and Experimental PlanTi-6Al-4V powder The material used for this investigation is EOS Ti64, which fulfills requirements of ASTM F1472 (for Ti-6Al-4V) regarding maximum concentration of impurities [9]. Fig.1 shows the powder morphology under a scanning electron microscope. The powder was measured using a Microtrac S3000 laser-based particle size analyzer. Its particles have a size distribution between 25 µm (D10) and 53 µm (D90). The particle size is nearly normally distributed with Mean Volume Diameter around 38 µm.50 µmFig. 1 EOS Ti-6Al-4V powder morphology and particle size distributionFatigue testing Fatigue specimens were built vertically along their axis with an as-built gauge surface in an EOS M270 Direct Metal Laser Sintering (DMLS) system.

The processing parameters from EOS are given in Table 1. Laser scanning along paralleled paths produced solidified layer. Contour scan was followed for each layer. It is believed that the specimen’s density is almost 100% of STA Ti64 material [7]. Specimens conform to the ASTM E466 standard with a continuous diameter between ends. Some specimens were designed with cylindrical or double-conical (D-conical) internal defect by leaving a void in the CAD file, as shown in Fig. 2.Table 1 Process parameters for fabricating Ti-6Al-4V fatigue specimens.Laser power (W) 170Scan speed (mm/s) 1250Hatch spacing (µm) 100Layer thickness (µm) 30Ø 0.5 mmØ 0.5 mm0.4 mmØ 3 mm0.8 mmØ 3 mm(a)(b)Fig. 2 CAD model of fatigue specimen (a) with cylindrical defect, (b) with D-conical defect500All specimens were heat treated at 650 Celsius for 4 hours in an argon filled furnace for stress relief and cooled down in the furnace to the room temperature.

Light sand blasting was used to clean adhered particles from the surface of the specimens. Fatigue tests were conducted on an Instron ElectroPuls E10000 All-Electric Test Instrument. The specimens were tested using a continuous sinusoidal load with frequency 50 Hz. Stress ratio R (min/max stress amplitudes) is 0.1. The fracture surfaces were then examined by optical microscopy (Olympus MX51 industrial inspection microscope) and scanning electron microscopy (FEI Nova NanoSEM 600) for analyzing the effect of defect on the result of fatigue tests.Results and DiscussionCharacteristics of Fatigue Specimens An as-built Ti-6Al-4V fatigue specimen is shown in Fig. 2. Through the scanning electron micrograph, it can be seen that the gauge surface is rough and irregular. Adhered particles were removed from the surface of the specimens by sand blasting. But some partially melted particles may be welded to the surface, resulting in convex surfaces which seem like speckles after blasting.

But some partially melted particles may be welded to the surface, resulting in convex surfaces which seem like speckles after blasting. But, in general, sand blasting does not affect the feature of as-built surface.Melting surfaceSpeckleFig. 3 As-built Ti-6Al-4V fatigue specimen and its surface featureFatigue results High cycle fatigue tests can cause cracking in fatigue specimens. Stress-life plots of the fatigue specimens are shown in Fig. 4. The stress life of fatigue specimens without designed defects was rather low at a stress level higher than 500 MPa. Three specimens were tested at 500 MPa. One of them failed very quickly, while another two did not fail at all. When the stress level was lowered to less than 500 MPa, even at 10 million stress cycles the specimens did not fail.It was assumed that the fatigue specimens would have lower stress life if a cylindrical or double- conical defect is included.

Therefore, all the specimens with internal defects were tested with a maximum stress no larger than 500 MPa in order to investigate how fatigue life is affected. The results show that the fatigue life at some specific stress level is repeatable, such as fatigue specimens with a cylindrical or double-conical defect at 450 MPa. But, at other stress levels, the fatigue life could be distributed across a range. For example, the fatigue life at a 500 MPa stress level is not repeatable, with values between sixty thousand and one million cycles, with no501discernible pattern as to cylindrical defects or double-conical defects. Thus it is hard to estimate the number of cycles to failure for the fatigue specimens at this stress level. When lowered to a 400 MPa stress level, the fatigue life of one specimen with a cylindrical defect was even twenty times more than two other specimens.

When lowered to a 400 MPa stress level, the fatigue life of one specimen with a cylindrical defect was even twenty times more than two other specimens. When the maximum stress was lowered to 350 MPa, fatigue specimens with defects did not fail until after ten million cycles.650Max Stress (MPa)600550500450400350Crack initiation from surfaceCrack initiation from centerWithout designed defect Not failed With D-conical defect Not failed With cylindrical defect Not failed Regression only for cracking from D-conical defect Regression only for cracking from cylindrical defect30010000100000100000010000000Cycles to failureFig. 4 Stress-life plot for fatigue testing of Ti64 specimens with and without geometric defectsFractography Generally the process of fatigue failure is characterized by crack initiation, crack propagation and final failure. Cracks usually initiate at the site with the highest stress concentration.

Cracks usually initiate at the site with the highest stress concentration. A microscopic view of Ti64 fatigue specimens shows a series of radiating ridges on the fracture surface, which can be traced along the direction to identify the area of the fracture initiation. These radiating ridges occupy most of the fracture surface. In the final fast fracture area, “cup and cone” fracture is formed on the matching surfaces oriented at about 45° from the tensile axis.For the fatigue specimens without designed defects, it was found that fatigue crack initiation only occurred at the surface. Fatigue fracture for some specimens occurred outside the gauge section and the fracture extended to the whole cross section and finally failed. Fig. 5 shows a typical fracture surface whose fatigue crack initiated at the gauge surface. A boundary deficiency can be seen at the crack initiation site. The crack propagated from the initiation site over several cycles.

A boundary deficiency can be seen at the crack initiation site. The crack propagated from the initiation site over several cycles. Final fracture occurred when the remaining area was too small to carry the maximum load of the final cycle. XRD analysis indicates as-built SLM Ti-6Al-4V contains an HCP phase with both α-phase and α’ martensite [10, 11]. Therefore, the radiating ridges refer to transgranular cleavage of α and α’ grains. The low failure cycles of these specimens can be attributed to crack growth from some of the α or α’ grains with very little crack initiation time. Micro striations on grain surfaces can be observed in the crack propagation region at a higher magnification as shown in Fig. 5 (e) and (f). It is assumed that a transition from transgranular cleavage to classic fatigue striations occurred and then changed to final fast fracture in ductile dimple mode [12]. The orientation of these striations varied from grain to grain.

The orientation of these striations varied from grain to grain. Dimpled appearance which is characteristic of microvoid coalescence is exhibited in the final fast fracture area. Some regions are located between dimpled areas and appear to be smooth at low magnification.502Crack initiationCrack propagationFinal fracture(a)(b)(c)(d)Fatigue StriationFatigue Striation(e)20 µm(f)Fig.5. Fractography of fatigue specimen (crack initiated from gauge surface) (a) Top view, (b) Crack initiation site, (c) Crack propagation (transgranular fracture), (d) Final fast fracture region, (e) Fatigue striation in crack propagation region (SEM), (f) Fatigue striation (optical microscopy)The fracture pattern of fatigue specimens with designed defects is varied, initiating either from the surface of the gauge section or from the designed defect in the interior. Thus, the stress-life plot can be divided into two areas for failed fatigue specimens as shown in Fig. 4.

Thus, the stress-life plot can be divided into two areas for failed fatigue specimens as shown in Fig. 4. The former503fracture pattern has a similar fractography as shown in Fig. 5, accompanied with a lower fatigue life than the latter fracture pattern. This explains why there is such large scatters amongst the fatigue life data for certain stress levels.When the fatigue crack was initiated from the designed cylindrical defect, the fracture surface is close to an edge of the defect, as shown in Fig. 6. The crack initiated from one of the defect’s boundaries. This indicates that geometry-dependent stress concentrations at the edge of a defect can be the cause of crack initiation. However, it is difficult to identify the specific crack initiation site with respect to one of these the α or α’ grain facets. The primary grains in the initiation site may all have orientation pointing along crack propagation.

The primary grains in the initiation site may all have orientation pointing along crack propagation. A transgranular cleavage took place in the crack propagation region from and around the crack initiation site, followed by ductile- dimple fast fracture.Crack initiationCrack propagationFinal fractureFig. 6 Fracture surface of fatigue specimen (crack initiated from central cylindrical defect)Fig. 7 shows a typical fracture surface with a crack initiated from the double-conical defect. Transgranular cleavage is the primary crack propagation pattern. Ductile dimple appearance is seen in the fast fracture region.Crack propagationCrack initiationFinal fractureFig.

Transgranular cleavage is the primary crack propagation pattern. Ductile dimple appearance is seen in the fast fracture region.Crack propagationCrack initiationFinal fractureFig. 7 Fracture surface of fatigue specimen (crack initiated from central double-conical defect)504According to the fundamentals of fracture mechanics, both cylindrical and double-conical defects should have a detrimental effect on the fracture strength because the applied stress will be amplified near the defect. The magnitude of this amplification depends on crack orientation and geometry. However, fractography shows no apparent difference in the fracturing process between cylindrical defects and double-conical defects. Moreover, the regression lines show similar tendency between these two types of defects as shown in Fig. 4, if only crack initiations from designed defect are considered in the regression.

Moreover, the regression lines show similar tendency between these two types of defects as shown in Fig. 4, if only crack initiations from designed defect are considered in the regression. This indicates that these two defect morphologies do not result in apparent difference of stress amplification with respect to crack initiation.The rough surface of the gauge section contains a random distribution of defects within the gauge surface. These defects negatively impact the fatigue life as stress risers and cause quick failure of Ti-6Al-4V fatigue specimens under the higher stress intensities. This is demonstrated from the stress-life plot in Fig. 4. It was found that as-built fatigue specimens are susceptible to fatigue crack initiation from the surface when the maximum stress is larger than 500 MPa. Fatigue specimens with designed defects should encounter the same situation, although no fatigue tests were conducted at stress levels above 500 MPa for those specimens.

Fatigue specimens with designed defects should encounter the same situation, although no fatigue tests were conducted at stress levels above 500 MPa for those specimens. At stresses below 500MPa, as-built surface defects have no apparent impact on fatigue life. This could be the result of microstructural characteristics around surface defects compared to those near designed internal defects. When Ti-6Al-4V fatigue specimens are free from any internal/surface defects or the defects are not sufficient enough to act as stress raisers then the fatigue properties are strongly influenced by the morphology and arrangement of the α and β phases. SLM fabricated Ti-6Al-4V exhibits a lamellar microstructure. In lamellar microstructures fatigue cracks initiate at slips bands within α (α’) lamellae or at α (α’) along prior β grain boundaries [13]. This can result in crack initiation to occur from the interior.

This can result in crack initiation to occur from the interior. Thus, in addition to the role of defects themselves, further investigations are merited to better understand the microstructural differences between skin and core areas of SLM-produced Ti-6Al-4V specimens to help understand how these might affect crack initiation.ConclusionDefects play a critical role in SLM Ti-6Al-4V fatigue specimen performance. As-built surfaces become crack initiation sites, especially when the stress intensity is higher than 500 MPa. Designed internal defects also affect the fatigue life of SLM Ti-6Al-4V materials so that the fatigue limit is lowered during fatigue testing. In these tests, no significant differences between cylindrical and double-conical defects were observed.

In these tests, no significant differences between cylindrical and double-conical defects were observed. In order to explain the effect of defects on Ti-6Al-4V fatigue life, further studies concerning the microstructure and morphology of the surface of fatigue specimens are required.AcknowledgeThe authors gratefully acknowledge the support of the Office of Naval Research, awards N00014‐09‐1‐0147 and N00014‐10‐1‐0800, Technical Monitor: Dr. Ignacio Perez. The authors also express their gratitude to the staff of the Rapid Prototyping Center at the University of Louisville for their assistance.505Reference[1] American Society for Testing and Materials (ASTM). Standard Terminology for Additive Manufacturing Technologies (F2792-12). http://www.astm.org.[2] I. Gibson, D.W. Rosen, B. Stucker. AdditiveManufacturing Technologies: Rapid Prototyping to DirectDigital Manufacturing. Springer, New York, NY, 2009[3] G. Levy, R. Schindel, J.P. Kruth.

viewed PaperEFFECT OF DEFECTS ON THE MECHANICAL PROPERTIES OF LASER POWDER BED FUSED Ti-6Al-4VMuztahid Muhammad1,2, Shuai Shao1,2, Nima Shamsaei1,2*1 National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849, USA 2 Department of Mechanical Engineering, Auburn University, Auburn, AL 36849, USACorresponding author: shamsaei@auburn.edu Phone: (334) 844-4839AbstractProcess-induced volumetric defects are inherent to additively manufactured parts. This study investigates the effect of volumetric defects on the tensile properties of the laser powder bed fused (L- PBF) Ti-6Al-4V specimens fabricated with large variations in process parameters (a total of six sets of process parameters). Cylindrical rods of L-PBF Ti-6Al-4V specimens were stress-relieved before removal from plates and machined to tensile specimens.

Cylindrical rods of L-PBF Ti-6Al-4V specimens were stress-relieved before removal from plates and machined to tensile specimens. The defect distribution of specimens resulting from each set of process parameters was analyzed using a high-resolution X-ray computed tomography machine. Quasi-static tensile tests were performed at room temperature using a servo-hydraulic MTS machine. Tensile results were correlated with defect statistics. No apparent difference was observed in the yield strength of the L-PBF Ti-6Al-4V specimens despite the large variations in the process parameters resulting in significant differences in defect content.

However, a considerable drop in ductility was observed for the specimens fabricated with insufficient energy.Keywords: Additive manufacturing, Laser powder bed fusion (L-PBF/LB-PBF), Ti-6Al-4V, Defects,Tensile properties.IntroductionAdditive manufacturing (AM), a layer-wise advance manufacturing technology, has received significant attention as a suitable replacement for conventional subtractive manufacturing methods due to its capability to fabricate parts in near-net-shape conditions with little to no post-processing [1]. The rapid progress of different AM technologies and a better understanding of the process-structure-property relationship of additively manufactured (AM) parts have resulted in increased adoption of AM parts in engineering applications [2,3]. Among different materials that can be fabricated using AM, Ti-6Al-4V is one of the most prominent alloys used in most commercial Ti applications.

Among different materials that can be fabricated using AM, Ti-6Al-4V is one of the most prominent alloys used in most commercial Ti applications. Conventionally manufactured Ti-6Al-4V components have been widely used in aerospace, automotive, naval, and biomedical sectors [4–6]. The combination of high strength and light weight as well as the ability to sustain in highly corrosive environments and elevated temperatures made Ti-6Al-4V highly desirable in these applications [4,7].AM method, although it offers numerous benefits over conventional manufacturing methods, AM components often consists of process-induced volumetric defects. The type of defects form in AM parts mainly depend on the fabrication process parameters (e.g., laser power, scan velocity, hatch distance, layer561thickness, etc.).

The type of defects form in AM parts mainly depend on the fabrication process parameters (e.g., laser power, scan velocity, hatch distance, layer561thickness, etc.). The most common volumetric defects in AM parts are keyhole (KH) pores (usually form due to excessive energy or heat input [8]), lack-of-fusion (LoF) defects (form mostly due to insufficient energy input[9]), and gas-entrapped pores (form due to gas entrapment, supersaturation of metal gas within the melt-pool, etc. [10,11]). In addition, the formation of defects in AM parts may also be affected due to powder characteristics [1,12,13], the volume of parts [14], geometry and location of parts on the build plates [15], etc., even after using recommended process parameters. These process-induced defects can impact the mechanical performance of AM parts [16–18].

These process-induced defects can impact the mechanical performance of AM parts [16–18]. It has been reported that the LoF defects, due to their irregular morphology, can cause a significant impact on the tensile properties of laser powder bed fused (L-PBF) Ti-6Al-4V [19–22]. Hence, the impact of defects on the mechanical performance of AM Ti-6Al-4V parts needs to be well understood for the qualification and certification for critical load- bearing applications [23,24].The formation of volumetric defects and their impact on mechanical properties for AM Ti-6Al-4V parts manufactured via optimized process parameters has been well studied [20,25,34,26–33]. However, based on the studies available in the literature, there is a lack of understanding of how defects type and density impact the tensile properties of L-PBF Ti-6Al-4V.

However, based on the studies available in the literature, there is a lack of understanding of how defects type and density impact the tensile properties of L-PBF Ti-6Al-4V. Therefore, this paper investigates the tensile properties of L-PBF Ti-6Al-4V fabricated with large variations in process parameters.Experimental ProcedureTi-6Al-4V cylindrical rods were fabricated employing an EOS M290 machine with six different process parameter sets. Apart from the manufacturer recommended process parameter, five other sets of process parameters were used to induce different types of KH and LoF defects. Process parameters (apart from the recommended one) varied so that insufficient or excessive energy induced LoF or KH type defects.The list of the process parameters used in this study is shown in Table 1. Following fabrication, the rods were stress-relieved at 705°C for 1 hour with a heating rate of 5°C/minute in an electric furnace and furnace cooled before removing from the build plate.

Following fabrication, the rods were stress-relieved at 705°C for 1 hour with a heating rate of 5°C/minute in an electric furnace and furnace cooled before removing from the build plate. Following stress-relieving, the Ti-6Al-4V rods were machined to round tension test specimens according to the ASTM E8 standard [35] (see Figure 1).Table 1. Process parameters employed for fabricating L-PBF Ti-6Al-4V parts.Intended defectsRecommended KH KH LoF LoF LoFDesignationsRecom. KHa KHb LoFa LoFb LoFcPower, P (W)280 336 364 252 224 280Layer thickness (µm)30Scan velocity, V (mm/sec) 1200 840 960 1200 1200 1200Hatch distance, H (µm) 140 140 140 140 140 168Energy Density, E (J/mm3) 55.55 95.24 90.28 50.00 44.44 46.29The defect type and population of specimens fabricated using each process parameter set was analyzed by scanning the middle of the gage section (length of scanned section is ~ 5 mm) employing a Zeiss Xradia 620 Versa X-ray computed tomography (XCT) machine.

XCT scanning was performed at 160 KV voltage and 21 W power along with an X-ray filter to reduce the count of low-energy photons.562The voxel size for each scan was kept at ~ 5.5 µm for all conditions. After scanning, the projection images were reconstructed using the Zeiss Xradia software and post-processed using ImageJ software.Quasi-static tensile tests were conducted using MTS servo-hydraulic testing machine with a load cell capacity of 100 KN according to ASTM E8M standard [35]. Tensile testing was conducted at a strain rate of 0.001 mm/mm/s in displacement-controlled mode. A mechanical extensometer was attached to the specimen to record the strain values till fracture. At least three specimens were tested for each set of process parameters.Figure 1. Geometry of the round tension test specimen according to ASTM E8M standard [35].

At least three specimens were tested for each set of process parameters.Figure 1. Geometry of the round tension test specimen according to ASTM E8M standard [35]. All the dimensions are in mm.Results and DiscussionIn general, the unique thermal history during fabrication of the AM parts contributes to the formation of process-induced defects (e.g., pores, LoF, etc.). The defect content may increase when thermal energy deviates from the optimized process parameters. Energy density for L-PBF parts can be calculated from the following equation (1) [36]:𝐸 =𝑃𝑣ℎ𝑡(1)where E is the energy density, P is the laser power, v is the scanning velocity, h is the hatching distance, and t is the layer thickness.Typically, excessive thermal energy input results in the formation of KH defects [37]. KH defects form when deep cavities formed due to vapor recoil are “pinched off” at the bottom of the melt pool during solidification [37–39].

KH defects form when deep cavities formed due to vapor recoil are “pinched off” at the bottom of the melt pool during solidification [37–39]. On the other hand, LoF defects are irregular in shape and typically form when the insufficient thermal energy density input fails to complete the melting and binding of metallic powders563[28,40,41]. However, regardless of formation types, excessive volumetric defects are undesirable as they can negatively impact mechanical properties [16].Scanned volumes of L-PBF Ti-6Al-4V specimens fabricated using different process parameters are shown in Figure 2. The defect distributions in specimens fabricated using different process parameters are also shown in terms of defect counts per volume and defect volume fraction in Figure 3.

The defect distributions in specimens fabricated using different process parameters are also shown in terms of defect counts per volume and defect volume fraction in Figure 3. As seen in Figure 2 and Figure 3, specimens fabricated in manufacturer-recommended process parameters and KHb parameters (i.e., the specimen fabricated using a 30% increase in laser power and 20% decrease in scanning velocity) contained the lowest defect count and density. In addition, the KHa specimen (i.e., the specimen fabricated using a 20% increase in laser power and 30% decrease in scanning velocity, with defect volume fraction of 0.019%) had a slightly increased defect volume fraction (factor of ~6) compared to the recommended (0.003%) and KHb specimens (0.002%). On the other hand, the number of defects and defect volume fraction of LoF specimens are much higher than both recommended and KH specimens (defect volume fraction of LoFa, LoFb, and LoFc specimens are 0.056%, 0.129%, and 0.571%, respectively).

It is also observed in Figure 3(a) that LoFb and LoFc specimens contained higher number of large defects compared to the recommended, KHa, and KHb specimens. Hence, it is evident that the formation of defects is more sensitive to the reduction of energy input than the increase. Moreover, among LoF specimens, the specimen fabricated with increased hatching distance (i.e., LoFc) consists of the highest defect volume fraction (0.571%), which indicates that increasing hatching distance is more detrimental than decreasing laser power input for L-PBF Ti-6Al-4V specimens.Figure 2. XCT scanned volume of L-PBF Ti-6Al-4V specimens fabricated using different process parameters: (a) Recommended, (b) KHa, (c) KHb, (d) LoFa (e) LoFb, (f) LoFc.564Figure 3.

Column charts in Figure 4(b) include yield strength (Sy), ultimate tensile strength (Su), and percent elongation to failure (%EL) of the specimens. As seen in Figure 4, both KH specimens (i.e., KHa and KHb) exhibit slightly higher tensile strengths (~ 40 MPa higher Sy and Su compared to the recommended conditions) compared to other conditions. In contrast, all the LoF specimens (i.e., LoFa, LoFb, and LoFc) exhibited comparable tensile strengths (i.e., both Sy and Su) to the specimens fabricated with recommended parameters. In addition, recommended and KH specimens (both KHa and KHb) showed comparable ductility (i.e., %EL); whereas all the LoF specimens had slightly lower ductility (i.e.,565%EL). All the LoF specimens (i.e., LoFa, LoFb, and LoFc) had similar %EL, ~ 20% lower than recommended and KH ones..Figure 4.

All the LoF specimens (i.e., LoFa, LoFb, and LoFc) had similar %EL, ~ 20% lower than recommended and KH ones..Figure 4. (a) Engineering stress-engineering strain diagram, and (b) column chart comparing tensile properties of L-PBF and wrought (collected from [42]) Ti-6Al-4V specimens fabricated using different process parameters.Comparable tensile strengths (i.e., both Sy and Su) of KH specimens compared to the recommended ones can be attributed to the nature of KH defects as well as thermal energy experienced by the KH specimens during the fabrication. It has been reported in the literature that KH defects, even if they form higher in number per volume, do not deteriorate the monotonic properties of AM parts to a significant extent [20]. In addition, excessive laser power can diminish the occurrence of any unmelted/partially melted powder particle in the specimens, thus reducing the possibility of having LoF defects [43].

In addition, excessive laser power can diminish the occurrence of any unmelted/partially melted powder particle in the specimens, thus reducing the possibility of having LoF defects [43]. On the566other hand, LoF specimens in this study were fabricated with 10%-20% less energy density compared to the recommended one resulting in 0.056%-0.571% defect volume fraction in the specimens, which had no apparent effect on the tensile strengths of the L-PBF Ti-6Al-4V specimens. However, the decrease (20%) in ductility (i.e., %EL) in LoF specimens in this study can be attributed to the higher fraction of volumetric defects compared to the recommended one. Volumetric defects acting as crack initiation sites of fracture contributed to the early onset of final fracture compared to the recommended one [20]. Voisin et al. [34] also reported that ductility (i.e., %EL) of AM Ti-6Al-4V is strongly correlated to the defect population present in the specimens.

Voisin et al. [34] also reported that ductility (i.e., %EL) of AM Ti-6Al-4V is strongly correlated to the defect population present in the specimens. Furthermore, all the L-PBF Ti-6Al-4V studied here exhibited higher tensile strengths (i.e., Sy and Su) but comparable ductility (i.e., %EL) compared to the wrought T-6Al-4V collected from MMPDS [42]. Based on the above discussion, it is evident that the monotonic test alone is not sufficient for the qualification and certification of AM parts; additional tests (e.g., fatigue) need to be conducted to confirm the quality of AM parts for critical load-bearing applications.Summary and ConclusionsThis study investigated the defect distribution and tensile properties of L-PBF Ti-6Al-4V fabricated using different process parameters. Tensile properties were correlated with defect types and density. The findings of this study can be summarized as follows:Upon variation of process parameters, defect distribution of L-PBF Ti-6Al-4V changed.

The findings of this study can be summarized as follows:Upon variation of process parameters, defect distribution of L-PBF Ti-6Al-4V changed. Insufficient energy density resulted in more defect contents than excessive energy density during the fabrication.L-PBF Ti-6Al-4V specimens even fabricated with defects using non-optimized process parameters exhibited comparable tensile strength to the recommended one.Defects in Ti-6Al-4V specimens negatively impacted the ductility compared to the recommended one, contributing to the early onset of the final fracture.AcknowledgmentThe authors greatly acknowledge the support provided by the Federal Aviation Administration (FAA) under Cooperative Agreement 12-C-AM-AU and the National Science Foundation (NSF) under grant # 1919818.References[1][2][3]J.W. Pegues, S. Shao, N. Shamsaei, N. Sanaei, A. Fatemi, D.H. Warner, P. Li, N.

fect of Different Graphite Materials on Electrical Conductivity and Flexural Strength of Bipolar Plates Fabricated by Selective Laser SinteringNannan Guo and Ming C. Leu Center for Aerospace Manufacturing Technology Department of Mechanical and Aerospace Engineering Missouri University of Science and Technology, Rolla, MO 65409AbstractGraphite is an excellent material for bipolar plates used in Proton Exchange Membrane (PEM) fuel cell due to its great chemical resistance, but the brittle nature makes it difficult to manufacture. Selective Laser Sintering (SLS) based on layer-by-layer manufacturing technology can fabricate graphite bipolar plates with complex gas flow channels. To improve the performance of bipolar plates including electrical conductivity and flexural strength, different graphite materials (natural graphite, synthetic graphite, carbon black, and carbon fiber) were investigated to fabricate test samples.

These samples then went through post processing including carbonization and infiltration. The results show that bipolar plates with electrical conductivity of 380 S/cm and flexural strength of 40 MPa are obtained from proper combinations of natural graphite and carbon fiber, which are higher than the target values set by the Department of Energy.1. IntroductionBipolar plate, which accounts for 40-50% cost and 60-80% weight of the whole fuel cell stack [1], is an important part in Proton Exchange Membrane (PEM) fuel cell assembly. The main functions of bipolar plate include carrying current away from each cell, distributing gas fuels within the cell and providing support for Membrane Electrode Assembly (MEA).

The main functions of bipolar plate include carrying current away from each cell, distributing gas fuels within the cell and providing support for Membrane Electrode Assembly (MEA). The Department of Energy proposed a technical target of bipolar plates for the year 2010 [2], in which the main requirements are electrical conductivity >100 S/cm and flexural strength >25 MPa.Compared with metal, graphite is a great material for bipolar plates due to its excellent chemical resistance and low weight. However, the brittle nature makes it difficult to manufacture. Recently, more and more researchers focus on graphite/polymer composite bipolar plate [3-7], which is easier to fabricate and has better mechanical strength but lower electrical conductivity compared to bipolar plates machined from bulk graphite.

Two main fabrication methods for graphite composite bipolar plates are injection molding and compression molding [4,5], which are suitable for mass production but less effective in the research and development stage, e.g. optimization of materials and design.Selective Laser Sintering (SLS), based on layer-by-layer manufacturing, has been applied to build graphite composite bipolar plates for PEM fuel cell [8-10]. The major advantages are its ability to build complex flow channels on the bipolar plates and the flexibility to investigate different materials and different designs. In this process, the mixture of graphite materials andbinder is scanned by laser and molten binder bonds graphite particles together to form 3D parts.

In this process, the mixture of graphite materials andbinder is scanned by laser and molten binder bonds graphite particles together to form 3D parts. Then the porous and weak green parts go through the post processes, carbonization and infiltration, to become gas impermeable, and to increase electrical conductivity and mechanical strength.One of the big issues of bipolar plates fabricated by SLS process is the relative low electrical conductivity [8]. In this paper, different graphite materials including Natural Graphite, Synthetic Graphite, Carbon Fiber and Carbon Black, were studied in order to obtain good electrical conductivity and mechanical strength, fulfilling the requirements for bipolar plates. All these materials and their mixtures were used to fabricate bipolar plate with SLS process. The properties were measured and compared with those fabricated by compression molding [11].2.

All these materials and their mixtures were used to fabricate bipolar plate with SLS process. The properties were measured and compared with those fabricated by compression molding [11].2. Materials and Processes2.1 MaterialsNatural graphite (3610), synthetic graphite (4437), carbon black (5303) and carbon fiber (AGM99) were obtained from Asbury Graphite Mills, Inc. (New Jersey, USA). The properties of these materials were shown in Table 1. Phenolic powder (GP-5546, Georgia Pacific) having typical size of 15µm was used as the binder.

(New Jersey, USA). The properties of these materials were shown in Table 1. Phenolic powder (GP-5546, Georgia Pacific) having typical size of 15µm was used as the binder. After carbonization, porous brown parts were infiltrated by liquid epoxy resin (EPONTM Resin 828) used as the matrix.Table 1 Properties of graphite materialsPropertiesNatural Graphite (3610)Size (µm) Surface Area (m2/g) Typical Conductivity (S/cm) *Provided by Asbury Graphite Mills, Inc., NJ, USA.75-1501.27 27.78Synthetic Graphite (4437)10-4511.46 17.24Carbon Black (5303)0.03254 2.93Carbon Fiber (AGM99) Diameter 7.4 Length 150 1.87 -2.2 Fabrication ProcessFirst, graphite powders and phenolic binder was ball milled for 12 hours. Then, an SLS machine (Sinterstation 2000) was used to build green parts. The parameters used for the SLS process were: fill laser power (12W), outline laser power (4W), laser scan speed (60in/s), layer thickness (0.004in), laser scan spacing (0.003in).

The parameters used for the SLS process were: fill laser power (12W), outline laser power (4W), laser scan speed (60in/s), layer thickness (0.004in), laser scan spacing (0.003in). The part bed and feed bins were maintained at 60oC and 40oC respectively. For carbonization, green parts were heated to 1000oC in the furnace filled with Argon gas, to convert the binder into carbon residue and get brown parts. The heating schedule was from room temperature to 200oC with a heating ramp rate of 60oC/h, followed by a slower rate 30oC/h to 600oC, and then a 50oC/h ramp rate to 1000oC, holding for 1h.

The heating schedule was from room temperature to 200oC with a heating ramp rate of 60oC/h, followed by a slower rate 30oC/h to 600oC, and then a 50oC/h ramp rate to 1000oC, holding for 1h. For infiltration, brown parts were immersed into the liquid epoxy resin, after 20min getting them out and putting into oven at 80oC for 1h to cure the resin.2.3 ExperimentsIn the experiments, Synthetic Graphite (SG), Carbon Black (CB) and Carbon Fiber (CF) were mixed with Natural Graphite (NG) respectively in different volume ratios, then mixed with 35vol% binder (in all experiments), which was based on the previous experimental result and can ensure that the green parts have enough mechanical strength to go through the processes that follow. These mixtures of graphite materials and binder were used to fabricate bipolar plates with the process mentioned above.

These mixtures of graphite materials and binder were used to fabricate bipolar plates with the process mentioned above. Then the properties of the final parts were measured.2.4 Characterization of PropertiesElectrical conductivity was measured by the Four Point Probe technique, following the ASTM C611 specification and using Keithley 2400 SourceMeter. Five 20×3×3mm3 specimens were tested and the average value was calculated. Flexural strength (dimension of samples: 3×10×60mm3) was measured with the three-point bending method, using Instron Model 4468. Microstructure of the samples was obtained by Hitachi S-4700 FE-SEM. Flowability of graphite powder was observed during the SLS building process.3. Results and Discussion3.1 Change of MicrostructuresDuring the whole fabrication process, the microstructure changed from green part to brown part and then to infiltrated part, as shown in Fig. 1.

Results and Discussion3.1 Change of MicrostructuresDuring the whole fabrication process, the microstructure changed from green part to brown part and then to infiltrated part, as shown in Fig. 1. Figure 1(a) shows the microstructure of a green part, in which the binder bonded graphite particles together. After carbonization, the binder was removed and lots of pores were left in the brown part (Fig. 1(b)). Finally these pores were filled with resin after infiltration and only few tiny voids were left (Fig. 1(c)).Fig. 1 Change of microstructures during the fabrication process: (a) green part; (b) brown part; (c) infiltrated part. The material composition is 15vol% SG, 50vol% NG and 35vol% binder.After infiltration, the flexural strength of bipolar plates increased from 1.56MPa to 38MPa, as shown in Table 2. The material used here was 65vol% SG and 35vol% binder.

The material used here was 65vol% SG and 35vol% binder. Additionally, the result of electrical conductivity shows that infiltration with resin had no distinct negative effect on electrical conductivity. This is because the good connection between conductive filler (graphite particles) had been established in the green parts and brown parts (Fig. 1(a), (b)), whichwas not broken by infiltration. This indicated that infiltration with liquid epoxy resin was a better way to increase the strength of fuel cell bipolar plate, meanwhile not impacting the electrical conductivity.Before infiltration After infiltrationTable 2 Properties of bipolar plates before and after infiltration Electrical conductivity (S/cm) 48.9 46.3Flexural strength (MPa) 1.56 38.4The material used was 65vol% SG and 35vol% binder.3.2 Effect of Synthetic GraphiteFigure 2 shows the microstructure of a brown part with SG and NG, in which the larger flaky particles are NG and the smaller sphere ones are SG.

Electrical conductivity with various volume fractions of SG was shown in Fig. 3 (keeping binder at 35vol%, with the rest being NG). When only NG (65vol%) and binder (35vol%) was used, the electrical conductivity was 380S/cm. As the SG content increased, the conductivity decreased because the conductivity of NG particles is higher than that of SG. When SG particles were introduced into NG powder, these particles hindered the conduction of current among NG particles. The results show that several values of conductivity were higher than the target value of DOE (100S/cm).Fig. 2 Microstructure of brown part with 15vol% SG, 50vol% NG and 35vol% binder.Fig. 3 Electrical conductivity varies with different SG fractions (keeping binder at 35vol%, with the rest being NG).The effect of SG on flexural strength of bipolar plates was shown in Fig. 4. SG has slightly negative effect on the flexural strength.

4. SG has slightly negative effect on the flexural strength. The strength decreased from about 37MPa to 33MPa when SG increased from 5vol% to 65vol%. Because the adding of smaller SG particles filled up the big voids among the larger flaky NG (as shown in Fig. 2), which reduced the porosity. Consequently a smaller amount of resin was absorbed after infiltration, thus less strength was established. But even so, all the results were still higher than the target value set by DOE (25MPa).Fig. 4 Flexural strength varies with different SG fraction.3.3 Effect of Carbon FiberCarbon fiber is widely used in composite materials to enhance mechanical strength. Figure 5 shows the microstructure of a brown part with CF and NG. It is shown that CF was uniformly mixed with NG particles.Fig. 5 Microstructure of a brown part with 25vol% CF, 40vol% NG and 35vol% binder.Electrical conductivity with different volume fractions of CF was shown in Fig. 6.

5 Microstructure of a brown part with 25vol% CF, 40vol% NG and 35vol% binder.Electrical conductivity with different volume fractions of CF was shown in Fig. 6. Electrical conductivity decreased as the fraction of CF increased. This is because the dispersion of CF among NG particles broke the contacts of NG particles and increased the electrical resistance.Figure 7 shows that flexural strength varies with different CF fractions. Flexural strength increased greatly (35MPa to 40MPa) after introducing CF, and kept increasing with increase in the volume fraction of CF. When the content of CF was 25vol%, the flexural strength was almost 50MPa, which was 1.5 times of that without CF. It is obvious that the presence of CF can increase mechanical strength greatly.Fig. 6 Electrical conductivity varies with different carbon fiber fractions.Fig.

It is obvious that the presence of CF can increase mechanical strength greatly.Fig. 6 Electrical conductivity varies with different carbon fiber fractions.Fig. 7 Flexural strength varies with different carbon fiber fractions.3.4 Effect of Carbon BlackIt was reported that nano-size CB could be dispersed among NG particles to increase electrical conductivity for graphite composite bipolar plates made by compression molding [11]. Chen, et al. [10] also used epoxy resin containing carbon black to infiltrate bipolar plates made by the SLS process, to improve electrical conductivity. In this study, CB was directly mixed with NG to make bipolar plates with SLS. Figure 8 shows the microstructure of a brown part made from 16vol% CB, 49vol% NG and 35vol% binder. As shown in Fig. 8(b), the surface of NG particle was covered by these nano-size CB particles.The variation of electrical conductivity with different volume fractions of CB was shown in Fig. 9.

As shown in Fig. 8(b), the surface of NG particle was covered by these nano-size CB particles.The variation of electrical conductivity with different volume fractions of CB was shown in Fig. 9. As the CB content increased, conductivity gradually decreased. After the fraction of CB reached 5vol%, electrical conductivity fell below 100S/cm. This is because CB has lower conductivity compared with NG. During the ball-milling process, CB particles with large surface area tended to agglomerate and cover on the whole surface of NG particles (as shown in Fig. 8), which hindered the contacts between NG particles. This is different from the case in compression molding [11], in which NG and CB were first mixed in liquid resin. CB was dispersed well and filled the small voids between NG particles rather than covered their surface, thus can enhance electrical conductivity.Fig. 8 (a) Microstructure of a brown part with NG and CB; (b) detail view of the surface of NG particle.

8 (a) Microstructure of a brown part with NG and CB; (b) detail view of the surface of NG particle. (16vol% CB, 49vol% NG and 35vol% binder).Figure 10 shows that flexural strength varies with different CB fraction. A similar result with SG was obtained in the case of adding CB into NG. The flexural strength slightly decreased with increase in the fraction of CB. A possible reason is that the coverage of CB on the surface of NG made it difficult for liquid resin (matrix) to fully wet the surface of NG particles (major filler), hence after the cure of resin the interface between the filler and matrix was impacted and the mechanical strength declined.Fig. 9 Electrical conductivity varies with CB fractions.Fig. 10 Flexural strength varies with different CB fractions.3.5 FlowabilityAs a powder based process, SLS requires powder to have good flowability in order to successfully build parts layer-by-layer.

10 Flexural strength varies with different CB fractions.3.5 FlowabilityAs a powder based process, SLS requires powder to have good flowability in order to successfully build parts layer-by-layer. During the process after one layer is scanned, a new layer of powder is spread on the surface by a roller. Therefore, if the powder does not have a good flowability, the original position of the former layer will be moved by the new layer, due to the large friction between these layers.Based on the experimental results, SG powder had better flowability than NG and CF powder, because the latter two have large size and irregular particle shape. For the mixture of NG and CB, the coverage of CB powder on NG surface increased the friction among the NG particles and also made it difficult to build parts with SLS.

For the mixture of NG and CB, the coverage of CB powder on NG surface increased the friction among the NG particles and also made it difficult to build parts with SLS. Hence in order to fabricate bipolar plates with complex and small channels, adding SG particles in the mixture was necessary in terms of improvement of flowability.3.6 Comparison with Compression MoldingFor electrical conductivity, the highest value of about 300S/cm was obtained from compression molding [11], and about 380 S/cm from the SLS process. The main reason is that for compression molding, graphite particles (conductive filler) were mixed with liquid resin (insulated matrix) at first. Thus the surface of graphite was covered by resin, reducing conductivity.

Thus the surface of graphite was covered by resin, reducing conductivity. However for the SLS process, only graphite powders were mixed together first and then green parts and brown parts were built in which all the particles contacted each other directly even after infiltration with liquid resin, so the conductivity was a little higher than that obtained from compression molding. For flexural strength, the value obtained from compression molding was around 45MPa [11] and from SLS was around 40MPa. So the SLS result was comparable to the compression molding result. This is because both of the strengths are mainly provided by solidified resin, which is hardly different with and without pressure.Sample of bipolar plates fabricated using the process discussed above, was shown in Fig. 11. The material composition used was 45vol% NG, 10vol% CF, 10vol% SG and 35vol% binder. Electrical conductivity was about 120S/cm, and flexural strength was 40MPa.

11. The material composition used was 45vol% NG, 10vol% CF, 10vol% SG and 35vol% binder. Electrical conductivity was about 120S/cm, and flexural strength was 40MPa. The feature dimensions are: active area 50×50mm2, thickness 3mm, channel width 1.5mm and depth 1mm.Fig. 11 Bipolar plate with a serpentine flow field. (active area 50×50mm2, thickness 4mm, channel width 1.5mm and depth 1mm)4. ConclusionsThe SLS process was successfully applied to fabricate bipolar plates. Different graphite materials, NG, CF, SG and CB, were investigated in this process, and the corresponding electrical conductivity and mechanical strength were measured. NG was better for electrical conductivity and CF can greatly increase the flexural strength, but both of them were not good for flowability. Nano-size CB, covering the surface of NG particles, had negative effect on both conductivity and strength.

Nano-size CB, covering the surface of NG particles, had negative effect on both conductivity and strength. Although adding SG decreased the conductivity and had slightly negative effect on strength, it was very good for flowability, which ensured the process had ability to build small and complex channels for bipolar plates. Finally, the material composition of 45vol% NG, 10vol% CF, 10vol% SG and 35vol% binder, was chosen to fabricate bipolar plates and the obtained properties satisfied the DOE requirements. The performance of the fabricated bipolar plates in a PEM fuel cell stack will be tested in the future.Acknowledgement: This project is supported by Air Force Research Laboratory under contract #FA8650-04-C- 5704.Reference:[1] Tsuchiya, H, Kobayashi, O. Mass production cost of PEM fuel cell by learning curve. Int. J.Hydrogen Energy, 2004, 29: 985-990.[2] http://www1.eere.energy.gov/hydrogenandfuelcells/mypp/pdfs/fuel_cells.pdf, website, 2007.

This eventually increases the lead time and the manufacturing cost. To address these challenges, this study explores the fabrication of gradient and laminar structures of titanium aluminide (Ti-Al) composite coatings deposited on Ti-6Al-4V substrate via a single step laser cladding (LC). The alterations in microstructural properties, chemical composition and phase analysis of the coatings reinforced with TiC were investigated as a function of laser energy density. Evaluation of the deposited samples reveals that FGM composite clads were fabricated from Ti-Al blended with TiC when LED was set at 17.50 J/mm2. At the selected LED, a thermo-positive reaction between the constituents’ materials was induced matrix phases) and it resulted in the formation of intermetallic compounds (e.g. Ti2AlC, with a microhardness more than that of the substrate (Ti-6Al-4V alloy).

Ti2AlC, with a microhardness more than that of the substrate (Ti-6Al-4V alloy). This study provides new insights on the selection of process parameters for the coating manufacturers while employing low cost- and time-effective LC process for fabricating functional graded Ti-Al coatings.and𝛼𝛼2𝛾𝛾Keywords: Functionally graded materials (FGM); Laser cladding (LC); Chemical composition; Titanium aluminide (Ti-Al).1. INTRODUCTION:Engineering components fabricated from functionally graded materials (FGM) are characterised with differing microstructural, mechanical and physical properties due to gradual alteration in chemical composition over its entire volume with a view to achieving superior service performance [1, 2]. FGMs ensure that its constituent dissimilar materials compensate for each other’s demerits so as to manufacture a product with improved service performance relative to each of the parent materials when fabricated singly.

Due to these reasons, FGMs are becoming increasingly popular in aerospace, nuclear, electronics, medical and chemical industries as seen in Nazarov and co-investigators [3] who demonstrated the potential of selective laser melting (SLM) fabricated functionally graded nickel aluminide (Ni-Al) system in manufacturing aviation components.1422FGMs have been manufactured from Fe3Al and SS316L [4]; Ti-6Al-4V and SS316 stainless [1]; nickel aluminide (Ni-Al) [3]; and titanium aluminide (Ti-Al) [5, 6] for various technological applications. Among various FGM systems, titanium aluminide (Ti-Al) are attracting increasing industrial and academic attention because of its superior elastic modulus, specific strength, creep strength and excellent temperature oxidation resistance [7, 8]. Nevertheless, Ti-Al structural parts are characterised with undesirable room temperature ductility, susceptibility to cracking and poor wear resistance [9-11].

Nevertheless, Ti-Al structural parts are characterised with undesirable room temperature ductility, susceptibility to cracking and poor wear resistance [9-11]. Meanwhile, it has been established that controlled dispersion of reinforced particulates into FGM titanium aluminide matrix could enhance its wear resistance as well as associated microstructural and mechanical properties [9].Among the manufacturing routes engaged in fabricating FGMs and its composites are selective laser melting (SLM) [3]; laser cladding (LC) [8, 9]; friction stir processing [12]; laser metal deposition [1, 4, 6] and self-propagating high temperature synthesis [5]. Of all these manufacturing routes, laser-based manufacturing (LbM) techniques such as LC imparts minimal heat-affected zone (HAZ) and dilution, reduced residual stress and improved metallurgical bonding when engaged in extending the service life of engineering components.

Therefore, this study employs LC combined with pre-heating in depositing Ti-Al reinforced particulate FGM composite on Ti-6Al-4V substrate. The use of pre-heating in eliminating crack-susceptibility of engineering parts had been demonstrated in LbM.Ti-6Al-4V is characterised with high strength and corrosion resistance, however, it’s technological application in the automotive, aerospace, and power-generation industries is hindered by its limited operating temperature (400 oC) and poor wear resistance [13, 14]. This challenge can be overcome by incorporating reinforced particles into the FGM Ti-Al matrix in an attempt to increase the range of service temperature at which parts made in Ti-6Al-4V can be utilised as well as improve its wear resistance. Meanwhile, a search through the databases reveals that only a single study [9] had dwelt on LC fabricated FGM reinforced particulate composites till date.

Meanwhile, a search through the databases reveals that only a single study [9] had dwelt on LC fabricated FGM reinforced particulate composites till date. The study by Abboud and co-investigators [9] only considered the effects of scan velocity, powder feed rates and number of layers on the geometrical characteristics, microstructural variation and wear properties of Ti-Al/TiB2 FGM composites. Findings from the study pointed out that the samples were porous and characterised with cracks.

Findings from the study pointed out that the samples were porous and characterised with cracks. In order to establish a basis for manufacturing FGM parts for extending the high temperature service life and wear resistance of Ti-6Al-4V components in the automotive, aerospace, and power-generation industries, this study advances on the work of Abboud et al., (1994) by considering the effects of laser energy density (LED) on the consolidation mechanism, microstructural evolution, microhardness, phase analysis and chemical composition of Ti-Al/TiC FGM composites fabricated with LC process combined with pre-heating. LED indicates the combined influence of laser power (P), scanning velocity (V) and beam diameter (d) in order to understand the role of thermal transfer in the solidification mechanism and microstructural development of Ti-Al/TiC FGM composite coatings.14232.

MATERIALS AND METHODOLOGY:2.1 Materials:This study employed commercially available elemental gas atomised aluminium (Al: 45- 90 µm), grade 1 titanium (Ti: 45-90 µm) and irregularly shaped titanium carbide (TiC: 45-90 µm) morphology. The purity of Ti, Al and TiC powders was 99.9%. The powders were supplied by TLS, Technik GmbH. The feed rate in g/min of each powder was determined via the flowability graphs. The component ratio of the FGM composite constituents was determined by using the ratio of the feed rate of the parent materials. Titanium aluminide (Ti-Al) system consisting of 50%wtTi- 50wt%Al was formulated initially and the morphologies of the blended powders and that of TiC are shown in Figures 1a and b respectively. Figures 1c and d show the EDS analysis identifying the elemental composition of the blended 50%wtTi-50wt%Al and TiC powders respectively.

of layer 5 4 3 2 1Composition of powder blends for each layer 60% Ti-Al + 40wt% TiC 70% Ti-Al + 30wt% TiC 80% Ti-Al + 20wt% TiC 90% Ti-Al + 10wt% TiC 100% Ti-Al2.2 MethodologyPrior to deposition of FGM composite materials, Ti-6Al-4V substrates with dimension 100 mm x 100 mm x 5 mm were cleaned with acetone in an attempt to remove impurities from its surface and enhance the quality of the laser clad. Table 2 shows the process parameters used in depositing the FGM reinforced particulate composite in this study. These parameters were adopted on the basis of a prior study carried out by Tlotleng et al., [10] in which it was established that Ti- Al parts with desirable mechanical properties can be fabricated within the range of selected parameters.

Laser energy density (LED) in J/mm2 is defined as shown in (i)𝐿𝐿𝐿𝐿𝐿𝐿 = 𝑃𝑃 � 𝑉𝑉 ∗ 𝑑𝑑(𝑖𝑖)Table 2: Process parameters used in depositing the FGM reinforced particulate composite.1.25, 1.50, 1.75 Laser power (kW) 1.50 Scan speed (m/min) 4.00 Spot diameter (mm) Energy density (J/mm2) 12.50, 15.00, 17.50 Pre-heating temperature (oC) 400.00 12.00 Standoff distance (mm) Shielding gas flow rate (l/min) 10.00 2.00 Carrier gas flow rate (l/min)The deposition of the FGM particulate reinforced composite in single track and multi- layers was achieved via a 3kW IPG continuous fiber laser system. In order to inject the FGM constituent materials into the melt pool on a substrate pre-heated at 400 oC, a 3-way coaxial nozzle system connected to a 5-axis CNC machine and a 1.5 bar GTV multi-hopper powder feed system was employed to serve this purpose. The powder feed system regulates the powder feed rate through its rotary speed.

The powder feed system regulates the powder feed rate through its rotary speed. As reported earlier on, the flowability graphs of the powders were used in determining the component ratio of the FGM constituent materials as shown in Table 1. The shielding and carrier gases used in the experimental work are both made of argon gas with the specified flow rates in Table 2. Samples produced in each parameter shown in Table 2 were replicated twice. Since the deposition is continuous at the middle of the tracks, cross-sections of the deposited FGM clad samples were obtained at the middle of the track for microstructural and1425compositional characterisation. Thereafter, the sectioned FGM clads were ground and polished to a 0.04 µm (OP-S suspension) surface finish with a Struers TegraForce-5 auto/manual polisher.

Thereafter, the sectioned FGM clads were ground and polished to a 0.04 µm (OP-S suspension) surface finish with a Struers TegraForce-5 auto/manual polisher. The samples were etched in Keller’s reagent for 2 to 3 minutes so that the microstructural characteristics were determined by Olympus optical microscope equipped with Analysis® software.A scanning electron microscope (SEM), Jeol JSM-7100F equipped with energy dispersive X-ray spectroscopy (EDS) was used to carry out the microstructural and elemental analysis of the FGM composites. The SEM-EDS facility equipped with a video camera for observing the sample stage as its height is controlled used the NSS software for analysis. The phase composition of the FGM composite clads was identified via X-ray diffraction (Model Bruker D8 Advance.) with a Cu Kα monochromator radiation source. Material PDF files were used in identifying the phases present in the FGM composite coatings.

Material PDF files were used in identifying the phases present in the FGM composite coatings. Microhardness values of the FGM composite coatings were determined by using a Matsuzawa Seiko Vickers microhardness tester model MHT-1. The indenting load used was 500 g with a dwell time of 10 seconds used for each indentation. Microhardness measurements were taken along the length of the coating at 100 µm spacing as well as the interface between the substrate and the FGM composite coating. The experimental findings reported in this study were the average values of the corresponding results of the two groups of experiments.3. RESULTS AND DISCUSSION:3.1 Analysis of optical micrographs of FGM composite cladsthe developed microstructures at its lower, middle and upper sections (see Figures 2a-i) were examined with a view to understanding how variation in LED has altered the microstructure across its volume.

Based on microstructural evidence obtained from Figure 2, irrespective of the amount of LED dissipated into the fabricating the FGM, the particle size of unmelted titanium carbide had diminished in varying degrees relative to the particle size of the starting powder which lies between 45 to 90 µm (Figure 1b). This suggests that TiC particles melted and dissolved in the melt pool. Furthermore, independent of the amount of LED employed during LC processing, it is also noted that the sizes of unmelted TiC particles in the FGM composite samples tend grow increasingly as follows: top layer bottom layer. This implies that increased consolidation occurred at the lower section in comparison to the top section of the FGM.

In concurrence with Shishkovsky and co-investigators [6], the growth in the size of unmelted TiC particles as well as increased consolidation from the top layer to the lower layer of the FGM composite samples, irrespective of the LC process parameters, could be attributed to the phenomenon of directional solidification in which heat is dissipated away from the top layer to bottom layer.Optical micrographs (OM) ofthe FGM composites revealingmiddle layer→→1426(a)(d)(g)Unmelted TiCAl matrixFine Dendrites & needlelike phases(b)(e)(h)(i)(f)(c) Figure 2: Optical micrographs of the FGM composite fabricated with 12.5 (a, b, c); 15.0 (d, e, f) and 17.5 J/mm2 (g, h, i) viewed at the top layer(a, d, g); middle layer (b, e, g) and lower layer (c, f and i).Figures 2a-c, d-f, and g-i show the optical micrograph images of the FGM composite fabricated with 12.50, 15.00 and 17.50 J/mm2 respectively.

These micrographs establish that the amount of TiC melting and dissolving in the melt pool varies as the LED dissipated during the LC processing of FGM samples. It is also evident from the micrographs that the sizes of unmelted 17.50 J/mm2. The nature of TiC particles reduce as follows: 12.50 J/mm2 microstructure evolving at the top, middle and bottom layers of FGM samples fabricated with LED → of 17.5 J/mm2 is presented in Figures 2g, h and i respectively. The sample made with LED of 17.50 J/mm2 is distinguished from those fabricated with 12.50 J/mm2 and 15.00 J/mm2 as its microstructure is characterised with unmelted TiC particles as well as fine dendrite and needlelike phases at all location across the volume of the FGM.

In addition, dendrites are finest at the top layer in comparison to the middle and lower layers.15.00 J/mm2→3.2 Scanning electron micrographs and EDS analysis of FGM composite cladsScanning electron micrographs (SEM) of the FGM composite samples fabricated with LED of 12.50, 15.00 and 17.50 J/mm2 are shown in Figures 3a-c, d-f, and g-i respectively with a view to further highlighting on the distinguishing characteristics of each sample. It is evident from Figures 3a-c that that the microstructures of the FGM composite consists of grey-coloured1427unmelted TiC particles (1), white-coloured ring region surrounding the unmelted TiC particle (2) and dark-coloured region (3). Comparative analysis of FGM composite samples fabricated with 12.50 J/mm2 (Figure 3a-c) and 15.00 J/mm2 (3d-f) reveals that the thickness of white-coloured ring regions are larger in the latter than the former.

The thickness of these white-coloured ring region is also noted to be increasing from top to lower layers of the FGM samples irrespective of the selected process parameters. Meanwhile, the width of regions 1 and 3 have reduced in samples fabricated with 15.00 J/mm2 (3d-f) relative to that of 12.50 J/mm2.EDS analysis was carried with a view to gaining insight into the chemical composition of the regions 1, 2 and 3 formed in FGM samples fabricated with 12.50 J/mm2 and 15.0 J/mm2. EDS findings reported in Table 3 confirm that the grey-coloured region 1 is rich in titanium with trace quantity of carbon; the white-coloured ring region 2 consists of aluminium, carbon and titanium; while the dark coloured region 3 consists of aluminium and carbon with trace quantity of titanium.

The formation of white-coloured ring region 2 indicates the occurrence of metallurgical reaction between aluminium, titanium and carbon while the existence of region 3 indicates there was diffusion of carbon to aluminium rich dark areas.(a)(d)(g)1132X32C(b)(e)(h)(i)(c) Figure 3: Scanning electron micrographs (SEM) of the FGM composite fabricated with 12.5 (a, b, c); 15.0 (d, e, f) and 17.5 J/mm2 (g, h, i) viewed at the top layer (a, d, g); middle layer (b, e, g) and lower layer (c, f and i).(f)1428BATable 3: Typical chemical composition of FGM samples fabricated with LED of 12.0 J/mm2 and 15.0 J/mm2.1 2 3C 0.87 8.33 5.17Al85.68 94.19Ti 99.13 5.98 0.64Analysis of Figures 3g-i reveal the evolution of distinct microstructural features for FGM composite samples fabricated with 17.5 J/mm2. For instance, three regions consisting of the core (A) surrounded by layers (B) and sandwiched between region C are shown in Figure 3h.

For instance, three regions consisting of the core (A) surrounded by layers (B) and sandwiched between region C are shown in Figure 3h. The area marked X in Figure 3h is a representative area consisting of A, B and C which is observed at higher magnification (Figure 4). Figure 4 reveals that the core region 1 is surrounded by layers 2, 3 and 4 with the layers sandwiched in lighter grey region 5.

The EDS analysis of the regions is shown in Table 4.Figure 4: SEM observation of regions A, B and C at higher magnificationTable 4: Typical chemical composition of regions A, B and C.1 2 3 4 5C 0.80 0.92 1.74 2.12 2.60Al0.97 11.65 20.88 37.56Ti 99.20 98.11 86.60 77.00 59.841429(a)60AlTiC6)% wt(555lA, i504T345240135002004006008001000Distance from the top of the clad layer (µm)12001400(b))% wt(6560AlTiC87lA556, iT5054453402351300200400600 1000 Distance from the top of the clad layer (µm)8001200140016000 1800(c))% wt(lA, iT7060504030AlTiC9 8 7 6 5 4 3 2 1 02002004006008001000Distance from the top of the clad layer (µm)1200Figure 5: EDS point test data for the composition curves of FGM composite fabricated with (a) 12.50 J/mm2 (b) 15.00 J/mm2 and (c) 17.50 J/mm21430)% wt(C)% wt(C)% wt(CAnalysis of Figure 4 and Table 4 confirm that an in-situ reaction synthesis occurred between Ti- Al and TiC powders when LED of 17.5 J/mm2 was dissipated for processing as isolated titanium rich (1) and aluminium rich (3) regions which are associated with FGM composites fabricated with lower LED (Figures 3a-c; d-f) have depleted in size (Figures 3g-i).

This occurred as aluminium and carbon atoms diffused inward into the titanium core.Figures 5a, b and c depict the EDS point test data for the composition curves of FGM composite fabricated with 12.50, 15.00 and 17.50 J/mm2 respectively. An observation of Figure 5 shows that the EDS of all FGM composite clads possess the following features in common: immediate lift and dip, as well as up and down stagger irrespective of the laser energy density employed during fabrication. Meanwhile, intersection is only evident across the length of the sample fabricated with 12.50 J/mm2 (Figure 5a) and at the top layer of the FGM composite clad deposited with 15.00 J/mm2 (Figure 5b). The occurrence of these features on the compositional curves suggests the trends in the elemental concentration across the volume of the FGM composite. It is evident from Figures 5a, b and c that at instances when titanium diffuses faster, both aluminium and carbon tend to diffuse very slowly and vice-versa.

It is evident from Figures 5a, b and c that at instances when titanium diffuses faster, both aluminium and carbon tend to diffuse very slowly and vice-versa. At 12.50 J/mm2, it is established that Ti atoms diffuse slower than aluminium and carbon atoms (Figure 5a). As the LED increases to 15.00 J/mm2 and 17.50 J/mm2, titanium atoms diffuse faster than aluminium and carbon atoms (Figures 5b and c). With LED set at 12.50 J/mm2, the concentration of diffused aluminium atoms are above the expected values of 50wt%, 45wt%, 40wt%, 35wt% and 30wt% for each of the 1st, 2nd, 3rd, 4th and 5th layers respectively whereas the concentration of dissolved titanium atoms across at all the layers are below the expected values. This could be attributed to the fact that the dissipated LED is only effective at melting much of aluminium atoms to the detriment of titanium atoms.

This could be attributed to the fact that the dissipated LED is only effective at melting much of aluminium atoms to the detriment of titanium atoms. Meanwhile, at LED of 15.00 J/mm2 and 17.50 J/mm2, the concentration of aluminium atoms are much well below the expected values across the layers while those of titanium are well above the expected concentration values across the layers. It is pertinent to note that the concentration of dissolved titanium atoms across the layers of the FGM composite clad fabricated with 17.50 J/mm2 are higher than those of 15.00 J/mm2 while the converse is true for aluminium. The concentration of diffused carbon atoms across layers is constant at values varying between 3.00 to 4.00 wt% for samples fabricated with 15.00 J/mm2 and 17.50 J/mm2. The concentration of diffused carbon atoms across FGM layers at 15.00 J/mm2 and 17.50 J/mm2 is lower than that of 12.50 J/mm2.

The concentration of diffused carbon atoms across FGM layers at 15.00 J/mm2 and 17.50 J/mm2 is lower than that of 12.50 J/mm2. Reduced concentration of aluminium below the expected values across the layers at 15.00 J/mm2 and 17.50 J/mm2 indicate that aluminium atoms might have evaporated as the specified LED might have induced processing temperature higher than the melting point of aluminium. Increased concentration of titanium atoms for samples deposited with 15.00 J/mm2 and 17.50 J/mm2 suggests increasing trend in the formation of intermetallic compounds between Ti, Al and C atoms.

Increased concentration of titanium atoms for samples deposited with 15.00 J/mm2 and 17.50 J/mm2 suggests increasing trend in the formation of intermetallic compounds between Ti, Al and C atoms. Meanwhile, it is anticipated that the phase composition of intermetallic compounds forming from metallurgical reactions between Ti, Al and C atoms will increase in the following order: 12.50 J/mm2P 15.00 J/mm217.50 J/mm2.→→14313.3 XRD analysis of FGM composite cladsXRD analysis (Figure 6 and Table 5) of the samples confirms the formation of Ti2AlC matrix phases while (titanium aluminide carbide) in the FGM composites in addition to some TiC phases are still retained within the microstructure. The relative intensity of diffractive in the FGM composite clads is significantly enhanced as LED increases. peaks (Figure 6) for matrix occurs as LED increases.

The relative intensity of diffractive in the FGM composite clads is significantly enhanced as LED increases. peaks (Figure 6) for matrix occurs as LED increases. This emphasizes that increase in the volume fraction of Findings reported for Figures 3 and 5 as well as Tables 3 and 4 support this claim that the propensity of metallurgical reactions involving Ti and Al tends to increase as LED increases. 2 at. pct. carbon at processing According to Liu & DuPont [15], for a melt pool containing 2200 oC, primary TiC1-x solidifies initially. Thereafter, a peritectic reaction in temperatures AlC occurs as the melt pool cools down to 2200 oC. Further cooling of which Ti the melt pool below 2200 oC leads to the precipitation of the ternary AlC because of depletion of carbon content in the liquid composition. Later, a different peritectic reaction in which L +and𝛼𝛼2𝛾𝛾𝛼𝛼2𝛼𝛼2>+ L >𝐶𝐶1−𝑥𝑥𝑇𝑇𝑖𝑖2→Ti(Al) occurs.

Later, a different peritectic reaction in which L +and𝛼𝛼2𝛾𝛾𝛼𝛼2𝛼𝛼2>+ L >𝐶𝐶1−𝑥𝑥𝑇𝑇𝑖𝑖2→Ti(Al) occurs. Liu & DuPont [15] also reported that the reaction L + AlC could occur at the final stage of LC processing. Given the available information from 𝑇𝑇𝑖𝑖2 𝛽𝛽 → 𝛼𝛼 the work of Liu & DuPont [15], the phase transformation scheme for the FGM composite clads is + 𝑇𝑇𝑖𝑖2 elucidated as follows: 𝑇𝑇𝑖𝑖2Ti(Al) occurs. Liu & DuPont [15] also reported that the reaction L + AlC could occur at the final stage of LC processing. Given the available information from 𝑇𝑇𝑖𝑖2 𝛽𝛽 → 𝛼𝛼 the work of Liu & DuPont [15], the phase transformation scheme for the FGM composite clads is + 𝑇𝑇𝑖𝑖2 elucidated as follows: TiC + L (TiAl) +Ti(Al) occurs. Liu & DuPont [15] also reported that the reaction L + AlC could occur at the final stage of LC processing.

Liu & DuPont [15] also reported that the reaction L + AlC could occur at the final stage of LC processing. Given the available information from 𝑇𝑇𝑖𝑖2 𝛽𝛽 → 𝛼𝛼 the work of Liu & DuPont [15], the phase transformation scheme for the FGM composite clads is + 𝑇𝑇𝑖𝑖2 elucidated as follows: AlCAlC𝑇𝑇𝑖𝑖2 in In agreement with Pedrix and co-investigators [16], increment in volume fraction of the FGM composite clads with increased LED as seen in the XRD analysis suggests that increased carbon in solid solution as LED increases leads to enhancement in the stability of elevated phase. Furthermore, the Ti-Al-C phase diagram also confirms that the solubility of temperature carbon in the phase [15]. Increased metallurgical reaction 𝛼𝛼 between Ti, Al and C at the highest LED adopted in this study (see Table 2) suggests that increased 𝛼𝛼2 laser power enhances available heat being dissipated to the powder blends.

Consequently, this raises the temperature of the melt pool as well as the propensity of the laser-material interaction to form an intermetallic compound as increasing amount of TiC dissolved into the melt pool. The precipitation of primary TiC from the melt upon solidification in all the samples could be attributed to the fact that the carbon content in the melt pool (Tables 3 and 4) is higher than 2 at. pct. irrespective of the LED adopted in fabricating the FGM composite clads.

pct. irrespective of the LED adopted in fabricating the FGM composite clads. Although, the dissolved carbon content in the melt pool varies depending on the LED adopted for processing.𝑇𝑇𝑖𝑖2𝑇𝑇𝑖𝑖2→ 𝛾𝛾→ 𝛼𝛼2(𝑇𝑇𝑖𝑖3→→𝛼𝛼→phase is higher relative to that of𝛾𝛾1432𝛼𝛼2Figure 6: XRD analysis of FGM composite clad samples fabricated with LED of 12.50, 15.00 and 17.50 J/mm2.Table 5: XRD analysis showing the phases presented in FGM composite samples fabricated with varying LED.Phase2 Theta AnglesPhase Composition12.50 J/mm2 15.00 J/mm217.50 J/mm2A: B: 𝛾𝛾 C: TiC 𝛼𝛼2 D: Ti2AlC E: Ti21.92,38.95,44.73,78.57 -TiAl -Ti3Al 35.65, 40.90,53.81 64.64, 77.9635.65, 40.90, 53.81, 76.92 40.91,53.91 35.75,36,06,40,50,53,60,55,46,71,25√ √ √ √√ √ √ √√ √ √ √ √Each of the FGM composite clads fabricated in this study consists of different five layers regarding the composition and structure of each layer.

Therefore, the composition and microstructure of each layer is a function of the composition of the powder blends as well as the LED. The dilution resulting from the melting of one layer upon the other as well as its solidification phenomenon is manipulated by the LED. It is expected that a minimum dilution between the first layer and the substrate is attained by selecting the appropriate LED parameter to impart the desirable composition. Further studies will explore the alteration in the composition across the five layers of the FGM composite clads with reference to the Ti-Al-C ternary system.14333.4 Microhardness of FGM composite clads) across the layers of FGM composite clad samples at varying LED from the top surface of the deposits to the substrate. It is clear from Figure 𝐻𝐻𝑉𝑉0.5 100 HV0.5) and significantly less than 7a that the microhardness across the layers was constant ( that of the substrate.

It is clear from Figure 𝐻𝐻𝑉𝑉0.5 100 HV0.5) and significantly less than 7a that the microhardness across the layers was constant ( that of the substrate. This outcome is not surprising as the applied LED of 12.50 J/mm2 is not effective to initiate metallurgical reactions between Ti, Al and C to ensure formation of intermetallics which could have imparted higher microhardness across the FGM layers. A marginal increase in microhardness across the FGM layers is noted for samples deposited with LED of 15.00 J/mm2 relative to that of 12.50 J/mm2. The microhardness across the layers is however less than that of the substrate. In similarity to that of the sample deposited with 12.50 J/mm2, there is no difference in the microhardness across the second to fifth layers except in the first layer in which the microhardness is 200 HV0.5. This indicates that the adopted process parameter is just sufficient to initiate gradient composition in the first layer alone.

This indicates that the adopted process parameter is just sufficient to initiate gradient composition in the first layer alone. With LED increasing to 17.50 J/mm2, the microhardness profile (Figure 7c) obtained across the FGM composite clad layers is like that reported for two-layer functionally gradient Ti-Al/TiB2 composite [9]. Except for the first layer, the microhardness across the last four layers is either equal to or higher than that of the substrate (Figure 7c). The impartation of microhardness values higher than that of the substrate at the last four layers of this sample points to (i) the Ti-Al matrix alloying with aluminium as well as the partial melting of TiC which results in the precipitation of primary TiC crystals; and (ii) refinement of the microstructure at these layers upon the occurrence of the phenomenon of rapid solidification.

Meanwhile, reduced microhardness reported for the sample at the first layer and the interface could be attributed to the undesirable influence of the directional solidification in which excessive heat being transferred from the top layers to the first layers and the interface might have induced microstructural defects at the first layer and the interface.Figure 7 shows the microhardness (~~Evidence from microstructural/microhardness studies, chemical composition and phase analysis of the samples indicates that reaction synthesis will not take place between Ti and Al when LED is less than 17.50 J/mm2. Therefore, the formation of intermetallic compounds (e.g. matrix phases) results from thermo-positive reactions between Ti-Al. Prior to Ti2AlC, the commencement of the reaction synthesis via LC processing, the powder blends were heated near or thereabout the melting of aluminium.

Prior to Ti2AlC, the commencement of the reaction synthesis via LC processing, the powder blends were heated near or thereabout the melting of aluminium. Consequently, molten aluminium reacts with Ti particles via a thermo-positive reaction which produces a range of products including the TiAl. As a result, the temperature of the whole system is following: 3Ti + Al deemed to rise abruptly and this induces a chain of reaction in the entire system.

and (c) 17.50 J/mm2.𝐻𝐻𝑉𝑉0.514354. Conclusions:1. Based on microstructural evidence obtained from this study, the particle size of unmelted titanium carbide diminished in varying degrees, depending on the applied LED, relative to the particle size of the starting powder which lies between 45 to 90 µm. This suggests that TiC particles melted and dissolved in the melt pool.2. During the LC processing of the FGM composite clads, it was discovered that when titanium diffuses faster, both aluminium and carbon tend to diffuse very slowly and vice-versa. Moreover, titanium diffuses faster than both aluminium and carbon at higher LED.3. When the LED was set at 12.50 J/mm2 and 15.00 J/mm2, the induced processing temperature associated with the synthesis reaction of the powder blends was just adequate to affect the melting of aluminium.4. FGM composite clads were fabricated from Ti-Al blended with TiC when LED was set at 17.50 J/mm2.

FGM composite clads were fabricated from Ti-Al blended with TiC when LED was set at 17.50 J/mm2. At the selected LED, a thermo-positive reaction between the constituents’ materials was matrix induced and it resulted in the formation of intermetallic compounds (e.g. Ti2AlC, phases). 𝛼𝛼25. The impartation of microhardness values higher than that of the substrate for the FGM composite clad fabricated with 17.50 J/mm2 could be attributed to (i) the Ti-Al matrix alloying with aluminium as well as the partial melting of TiC which results in the precipitation of primary TiC crystals; and (ii) refinement of the microstructure at these layers upon the occurrence of the phenomenon of rapid solidification.6. The composition and microstructure of each layer of the FGM composite is not only a function of the composition of the powder blends but also that of the LED.REFERENCES:L. Thijs, K. Kempen, J. P.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEFFECT OF ENVIRONMENTAL VARIABLES ON TI-64 AM SIMULATION RESULTSAaron Flood* and Frank Liou*Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractIn metal AM the environment is critical and therefore care should be taken to ensure that the simulation matches reality. This paper will investigate the effect that various environmental factors have on the results of the simulation. This will help to determine their importance in the simulation setup. The material properties which relate to this are the convective coefﬁcient and the emissivity. These material properties will be investigated to determine their effect on the outcome of the simulation.

These material properties will be investigated to determine their effect on the outcome of the simulation. In addition to these properties, the size of the substrate will be investigated to determine if any results are altered. Lastly, the ambient temperature will be investigated to determine the effect this has on the simulation results.IntroductionWith the expense that is associated with additive manufacturing (AM), especially metal AM, in recent history many mathematical models [1, 2, 3, 4] have been developed to facilitate faster and cheaper understanding of the process. In order to properly simulate the process of AM, several assumptions and shortcuts need to be taken in order for the simulation to complete in an acceptable amount of time. An example of the use of shortcuts is to use static material properties instead of temperature dependent properties. These assumptions and shortcuts can have a signiﬁcant impact on the results of the simulation.

These assumptions and shortcuts can have a signiﬁcant impact on the results of the simulation. This study investigates the effect of surface material properties, ambient temperature, and substrate size on the temperature distribution on the surface of the spec- imen. The software used to run the simulation is the Additive Manufacturing Simulator (AMS). This software is being developed in a joint effort between Missouri University of Science and Technology and Product Innovation and Engineering, LLC. This software is an efﬁcient physics based simulation solution focusing on metal additive manufacturing.For all of the simulations which are performed a base set of parameters was used. These are listed in Table 1. It was determined that parameters of emissivity and convective coefﬁcient are so closely coupled that they should be investigated together, whereas the parameters of ambient temperature and substrate size were independent and could be investigated on their own.

In order to quantify the differences in the simulation results, a contour plot of the surface temperature was created for each simulation, an example of this can be seen in Figure 1. These contour plots were then analyzed to determine the number of voxels which were above a range of temperatures. A range of temperatures was used in order to avoid a bias which could have occurred if only a one or two temperatures were used.

A range of temperatures was used in order to avoid a bias which could have occurred if only a one or two temperatures were used. These results were then processes using SAS in order to determine any statistical differences.186550075010001250150017502000500 K1500 K2000 K1250 K750 K1000 K1750 KTable 1: General parameters used for all simulationsMaterial Laser Diameter Laser Power Laser Distribution Laser Scan Speed Resolution Emissivity Convective Coefﬁcient Ambient Temperature Substrate SizeTi-64 3.0 mm 1000 W Top Hat 5 mm/sec 750 µm 0.4 W/m2 5.0 W/m2K 298 K 1.5 cm x 1.0 cm x 0.75 cmFigure 1: Sample contour plot of surface temperatureMaterial Properties EffectOne of the most important aspects to any simulation is gathering the correct inputs. In the case of AM simulation, the material properties can be some of the most important and challenging to ﬁnd.

In the case of AM simulation, the material properties can be some of the most important and challenging to ﬁnd. This set of simulations will focus on the effect of the convective coefﬁcient and emissivity on the temperature distribution of the surface of a part. The values which were used as the base values in the design of experiment (DOE) can be seen in Table 2. These values were found to be a good representation of the wide range of values that can be found throughout literature.In order to realize the goal of determining their effect, the values needed to be varied by a substantial amount. For the emissivity, the constraint of the value needing to be between 0 and 1 was included and for the convective coefﬁcient, the constraint of needing to be greater than or equal to 0 was added. With this constraint the values chosen can be seen in Table 3.

With this constraint the values chosen can be seen in Table 3. This range of parameters resulted in the need for 10 experimental runs to determine the effect that each has on1866Table 2: Values found in literatureValue Convective Coefﬁcient 5.0 [2] 0.4 [5]PropertyEmissivityTable 3: Values used in material property investigationProperty Emissivity Convective CoefﬁcientHigh Value Base Value Low Value 0.4 5.00.9 7.50.1 2.5the results. A sample of the temperature distribution which resulted from the simulations can be seen in Figure 2.(a) Simulation number 2(b) Simulation number 0Figure 2: Sample images from material properties simulationsFrom the images collected, the contours were produced and the raw data was collected and can be seen in Table 4. These results were then analyzed to determine the ability of these twoTable 4: Raw data collected for material property simulationSim. Num. Conv. Coef. (p-value) Emis.

The p-value of the model is reported in Table 5. As can be seen from the p values reported, the model is a good ﬁt for the data, meaning1867Table 5: Statistical model analysis of material property resultsTemperature Model P Value500 750 1000 1250 1500 1750 2000< .0001 < .0001 < .0001 < .0001 < .0001 < .0001 < .0001that the parameters of convective coefﬁcient and emissivity have a statistically signiﬁcant effect on the temperature distribution on the surface of the part. Next, the effect that each parameter has on the model was investigated. The results can be seen in Table 6. This statistical analysis agrees withTable 6: Statistical analysis of variables for material property experimentsTemperature Conv. Coef.

The results can be seen in Table 6. This statistical analysis agrees withTable 6: Statistical analysis of variables for material property experimentsTemperature Conv. Coef. Emis.500 750 1000 1250 1500 1750 2000< .0001 < .0001 < .0001 < .0001 < .0001 < .0001 < .00010 0 0 0 0 0 0the results which can be observed in the raw data in Table 4, namely that changing the emissivity does not affect the temperature distribution of the surface of the part. In the raw data, this can be illustrated with simulation runs 0 and 1. In these simulations, the convective coefﬁcient is kept constant and the emissivity is changed, however, the results are identical. In the statistical analysis, this inference can be drawn from the p-value of emissivity being 0.Ambient Temperature EffectTangentially related to the material properties is the ambient temperature of the build chamber.

In order to investigate the effect the ambient temperature has on the simulation results the 4 values, listed in Table 7, were used.Table 7: Ambient temperatures used in simulationsProperty TemperatureTemperature (in degree K) 298., 500., 1000., 1500.When the simulations were completed the results were collected and the images were visually compared, a representative pair of images are shown in Figure 3. The image in Figure 3a is using1868(a) Simulation Number 1 (500◦K)(b) Simulation Number 3 (1500◦K)Figure 3: Sample images from ambient temperature simulationsa 500◦K ambient temperature and Figure 3b is using a 1500◦K. Other than the obvious difference in the substrate temperature, there is a visible difference in the melted region shown in red. When the ambient temperature is larger the melt pool also grows wider. This is also observed in the raw data which was collected and displayed in Table 8.

When the ambient temperature is larger the melt pool also grows wider. This is also observed in the raw data which was collected and displayed in Table 8. As can be observed in this table, asTable 8: Raw data collected for ambient temperature simulationsSim. Num. Amb. Temp 0 1 2 3298.0 500.0 1000.0 1500.01750◦K 2000◦K 2250◦K 2500◦K 2750◦K 2950◦K 1584 1669 1669 21991391 1465 1463 16941344 1360 1432 15451294 1307 1327 13541284 1304 1317 13411284 1304 1317 1341the ambient temperature is increased so does the number of voxels which are counted for each threshold temperature. In order to quantify these results, statistical analysis was performed by using a T-test to determine the signiﬁcance of the ambient temperature on the melt pool size, these are displayed in Table 9.

In order to quantify these results, statistical analysis was performed by using a T-test to determine the signiﬁcance of the ambient temperature on the melt pool size, these are displayed in Table 9. As can be seen, the ambient temperature has a statistically signiﬁcantTable 9: Statistical analysis for ambient temperature simulationsTemperature Model P Value1750 2000 2250 2500 2750 29500.0011 0.0002 < 0.0001 < 0.0001 < 0.0001 < 0.0001impact on the temperature distribution of the part. Interestingly, as the investigation temperature increases so does the signiﬁcance that the ambient temperature has on the surface area that is above that temperature. This effect could be an artifact of the statistics, namely a lower signal to noise ratio. In order to show this, a more detailed study must be performed.1869Substrate Size EffectThe ﬁnal environmental variable investigated which could affect the simulation is the ﬁxturing which is used to clamp the specimen in place.

This set of simulations investigates this effect by increasing the size of the specimen. This will simulate the increased thermal mass which can be seen when additional ﬁxturing is needed. The substrate sizes simulated can be seen in Table 10. The x and y dimensions of the specimen were increased proportionally while keeping the z dimension constant. Representative images from the simulations can be seen in Figure 4. VisuallyTable 10: Substrate sizes used in simulationsy x 1.5 1.0 0.75 1.65 1.1 0.75 1.2 0.75 1.8zvolume 1.125 1.36125 1.62Values (in cm or cm3)(a) Simulation number 0(b) Simulation number 2Figure 4: Sample images from substrate size simulationsinspecting the images, it can be seen that more conduction takes place in the larger part, Figure 4b, and less conduction occurs in the smaller specimen, Figure 4a.

The raw data for this experiment, Table 11, shows a similar trend, namely that larger substrates results in more heat being pulled away from the melt pool resulting more voxels at higher temperatures. The exception to this trend is simulation number 1. More conduction is seen in simulation 1 than simulation number 2 which was a larger substrate. In order to understand this phenomenon, more investigation is necessary. To quantify the statistical signiﬁcance of the raw data, a T-test was performed on the data, whichTable 11: Raw data collected for substrate size simulations500◦K 1000◦K 1500◦K 2000◦K 2500◦K 1.125 2804 1.36125 4021 4120 1.62Sim. Num. Volume 0 1 21851 2810 28751539 2175 20841308 1670 15961304 1415 1368resulted in the data in Table 12.

Num. Volume 0 1 21851 2810 28751539 2175 20841308 1670 15961304 1415 1368resulted in the data in Table 12. As can be seen from the p-values, the model is a good predictor of1870Table 12: Statistical analysis for substrate size simulationsTemperature Model P Value500 750 1000 1250 1500 1750 20000.0132 0.0193 0.0169 0.0114 0.0104 0.0056 0.0052the temperature distribution within the specimen. An interesting trend that can be seen in the data is that as the temperature is increased the p-value decreases. This shows that the substrate size is more critical to the temperature distribution at higher temperatures.ConclusionSeveral conclusions can be drawn from the simulations which were performed in this presented work which correlate the sensitivity of the temperature distribution to various environmental pa- rameters. The ﬁrst investigated were the emissivity and convective coefﬁcient.

The ﬁrst investigated were the emissivity and convective coefﬁcient. It was seen that the simulation is sensitive to the input of the convective coefﬁcient and not emissivity. When investi- gating the effect of ambient temperature, it was found that the ambient temperature has a statistical inﬂuence on the temperature distribution within the specimen. And lastly, it was seen that the substrate size has a signiﬁcant impact on the temperature distribution of the specimen. During the course of these simulations, several new questions were raised.In order to in- vestigate these questions, several future steps have been identiﬁed. First, the simulations will be performed at a higher resolution. This will identify effects that are particular to this resolution selection and ones which can be applied generally to the simulation. Along the same line of in- vestigation, the software allows for a dynamic resolution which will allow for a more detailed investigation of the melt pool area.

Along the same line of in- vestigation, the software allows for a dynamic resolution which will allow for a more detailed investigation of the melt pool area. In addition, the effects that these parameters have on the melt pool depth would like to be investigated. Lastly, the effect that various parameters have would like to be experimentally validated to ensure that the simulation results match reality.AcknowledgmentsThe support from Department of Energy Grant DE-SC0015207 and NSF grants CMMI 1625736 and EEC 1004839 are appreciated. We also appreciate the ﬁnancial support provided by the Cen- ter for Aerospace Manufacturing Technology (CAMT) at the Missouri University of Science and Technology.References[1] M. F. Z¨ah and S. Lutzmann, “Modelling and simulation of electron beam melting,” ProductionEngineering, vol. 4, no. 1, pp. 15–23, 2010.1871[2] Z. Fan and F.

fect of Extrudate Swell, Nozzle Shape, and Convergence Zone on Fiber Orientation in Fused Deposition Modeling Nozzle FlowB. P. Heller*, D. E. Smith*, and D. A. Jack*Department of Mechanical Engineering, Baylor University, Waco, TX 76798AbstractRecent advances for improving the mechanical properties of materials used in Fused Deposition Modeling (FDM) include the addition of carbon fibers to the filament feedstock. During processing, the flow field within the polymer melt orients the fiber suspension, which is important to quantify since fiber orientation influences mechanical and thermal properties. This paper presents a computational approach for evaluating polymer melt flow and fiber orientation within a FDM nozzle taking into consideration the converging flow in the nozzle, fluid expansion caused by extrudate swell, and nozzle exit shape.

Finite elements are used to evaluate the Stoke’s flow in an axisymmetric nozzle and fiber orientation tensors are evaluated along streamlines within the flow using the Fast Exact Closure and Folgar-Tucker isotropic rotary diffusion. Fiber orientation is shown to increase in the shear-dominated flow through the nozzle, however, alignment is found to decrease in the expansion flow of the die swell.IntroductionFused Deposition Modeling (FDM) is quickly becoming an industrially viable Additive Manufacturing (AM) technology. Intricate three dimensional parts can be created using low cost materials with little waste and low energy consumption. Another benefit to FDM is the ability to scale the technology to print very large parts with high dimensional accuracy. There still exist several issues with FDM that need to be addressed, such as warping, delamination, and unsatisfactory mechanical properties of typical FDM virgin polymers.

There still exist several issues with FDM that need to be addressed, such as warping, delamination, and unsatisfactory mechanical properties of typical FDM virgin polymers. The addition of carbon fibers to a virgin polymer was shown to significantly improve material integrity issues in FDM [1]. The addition of carbon fibers increases the strength and toughness of printed parts as well as reducing the CTE of the polymer fiber composite matrix. Unfortunately, little research has gone into the study of fiber orientation in the FDM process. Nixon, et al. [3] performed an initial study of the fiber orientation in three different FDM nozzle configurations (straight, converging, and diverging) with Moldflow (Moldflow Corporation, Framingham, MA) which uses the Advani Tucker [2] orientation tensor approach. The model produced by Nixon, et al. has the fiber filled polymer being injected into a large open mold and free surface open air boundary conditions are not considered.

The model produced by Nixon, et al. has the fiber filled polymer being injected into a large open mold and free surface open air boundary conditions are not considered. The fiber orientation in the polymer melt flow beyond the nozzle exit where the fluid boundary is partially formed by a free surface has yet to be studied for the FDM process. The extruded polymer melt will undergo extrudate swell at the nozzle exit which will affect the fiber orientation in the extruded bead; therefore, it is necessary to model this phenomena to produce an accurate model.Fiber orientation in compression molding, injection molding, and extrusion has been studied for several decades and is well understood. The computationally efficient fiber orientation tensor approach introduced by Advani and Tucker [2] has become the standardmethod for fiber orientation calculation.

The computationally efficient fiber orientation tensor approach introduced by Advani and Tucker [2] has become the standardmethod for fiber orientation calculation. Fiber orientation software packages such as Moldflow (Moldflow Corporation, Framingham, MA) and Moldex3D (Core Tech Systems Co., Ltd., Chupei City, Taiwan) use the Advani and Tucker orientation tensor model to predict the fiber orientation in fiber filled molded parts. With respect to extrudate swell, studies have also been conducted on fiber orientation in rapid geometric changes (see [4] and [5]) which is beneficial to understanding how the fibers will react in the extrudate swell region. The inclusion of the free surface of the extrudate swell to the model will have an impact on the fiber orientation and therefore the material properties of the final printed part.The swell of an extrudate at the exit of an extrusion die has also been studied and the phenomena is well understood.

Extrudate swell will occur at the nozzle exit of the FDM nozzle and is therefore of interest to the current study. The extrudate swell has been modeled for Newtonian fluids by Georgiou, et al. [6] and Reddy and Tanner [7] using similar methods with slightly different parameters. The swell of the extrudate of a Newtonian fluid was calculated and observed to be approximately 13% from the nozzle exit diameter.This paper presents a computational approach for computing fiber orientation in an axisymmetric FDM extrusion nozzle and within the polymer melt flow immediately after the polymer melt passes through the nozzle exit. The nozzle shape, convergence zone, and extrudate swell, and the effect of each on the fiber orientation state will be the focus in the current study. Fiber length is assumed to be sufficiently small to not clog the nozzle during printing.

Fiber length is assumed to be sufficiently small to not clog the nozzle during printing. To model the fiber orientation we use the orientation tensor model of Advani-Tucker [2] and the Isotropic Rotary Diffusion model from Folgar and Tucker [8]. The Fast Exact Closure of Montgomery- Smith, et al. [9] is used to solve the orientation tensor closure problem. The fiber orientation is calculated for polymer melt flow through a 0.35 mm diameter FDM nozzle which is typical of many commercially available FDM devices, such as the Makerbot 2X (Makerbot Industries, LLC, Brooklyn, NY). The nozzle geometry is then changed in a parametric study to better understand the effect of nozzle geometry on fiber orientation.Modeling ApproachThe cross section of a common FDM extrusion nozzle is used to create the polymer melt flow domain.

An axisymmetric slice is taken from the cross section seen in Figure 1 and is used to define a finite element model of the flow domain in the COMSOL Multiphysics (Comsol, Inc., Burlington, MA) program as shown in Figure 2. Axial symmetry can be used as a computationally beneficial and physically accurate representation of a three dimensional flow. In the current study the analysis is restricted to axisymmetric flow in the nozzle and a region just outside of the nozzle exit. The boundary conditions used in the model appear in Figure 2.It is assumed that the flow entering the nozzle is steady state laminar flow and thus the inletvelocity 𝑣𝑧𝑖𝑛𝑙𝑒𝑡 in the z direction at z = 5.625 mm is defined as𝑣𝑧𝑖𝑛𝑙𝑒𝑡 = −𝑣𝑚𝑎𝑥 [1 − (𝑟𝑖𝑛𝑙𝑒𝑡 𝑟𝑚𝑎𝑥2 )]with vx = vy = 0 at the inlet. In Equation (1), 𝑣𝑚𝑎𝑥 is the maximum velocity at the inlet of the fluid domain, 𝑟𝑖𝑛𝑙𝑒𝑡 is the radial location, and 𝑟𝑚𝑎𝑥 is the maximum radius of the fluid domain at and 𝑟𝑚𝑎𝑥 = 0.875 𝑚𝑚.

In Equation (1), 𝑣𝑚𝑎𝑥 is the maximum velocity at the inlet of the fluid domain, 𝑟𝑖𝑛𝑙𝑒𝑡 is the radial location, and 𝑟𝑚𝑎𝑥 is the maximum radius of the fluid domain at and 𝑟𝑚𝑎𝑥 = 0.875 𝑚𝑚. This inlet velocity allows the inlet. For this study we use 𝑣𝑚𝑎𝑥 = 4𝑚𝑚𝑠(1)for an average nozzle exit velocity of 65 most desktop FDM printers.𝑚𝑚𝑠which is within the range of extrusion speeds forFigure 1. Cross Section of FDM Extrusion Nozzle, Dimensions are in millimeters (mm)Figure 2. Boundary Conditions and Mesh for Nozzle DomainThe outlet at z = 0 mm is defined as 0 Pa pressure. The exterior wall (right-hand side in Figure 2) of the fluid domain is defined as a no-slip wall boundary with 𝑣𝑧 and 𝑣𝑟 both equal to zero for z  0.875 mm. A slip wall condition is defined along the free surface defining the extrudate swell boundary for z < 0.875 mm.. A slip wall is defined by COMSOL Multiphysics asn·v = 0 where n is the surface normal and v is the velocity vector.

A slip wall is defined by COMSOL Multiphysics asn·v = 0 where n is the surface normal and v is the velocity vector. The edge of the model at r = 0 mm, an axisymmetric boundary condition is imposed.The incompressible Newtonian fluid properties used in this simulation are density ρ = 1040 kg/m3 and dynamic viscosity of μ = 350 Pa∙s. These values are chosen since they are for an ABS plastic, which is a common FDM printing material used in industry, at 230 degrees C and under a shear rate of 575 s-1, a value that is typical in the FDM nozzle flow domain. It is understood that ABS, is a shear-thinning non-Newtonian fluid. Assuming an isothermal Newtonian fluid simplifies the simulation process and is expected to be a good first step into the study of fiber orientation in deposition processes.

Assuming an isothermal Newtonian fluid simplifies the simulation process and is expected to be a good first step into the study of fiber orientation in deposition processes. Use of a Newtonian fluid also makes it possible to compare die swell results to known data from Georgiou [6] and Reddy and Tanner [7].The finite element model is defined using the boundary conditions shown in Figure 2. A small sample of the mesh used is shown in Figure 2 as well. A mesh convergence study was performed during our preliminary studies to insure that a sufficiently fine mesh was employed so that the discretization would not affect the final results. The fluid flow problem is evaluated in Comsol Multiphysics using LiveLink which employs Matlab (The Mathworks, Inc., Natick, MA). The optimization approach described in Heller [?] which minimizes the combined shear and normal stress along the free surface is used here to compute the shape of the flow domain immediately outside of the nozzle exit.

These results predicted a radial die swell of approximately 13% for straight tube flow which agrees well with values obtained by Georgiou, et al. [6] and Reddy and Tanner [7].Fluid velocity and velocity gradients are calculated over the fluid domain, and fiber orientation through the nozzle is computed along streamlines within the flow domain. Ten streamlines appearing in Figures 4, 8, 9, and 11 (a) are computed at equally spaced locations across the nozzle outlet. At every point along the streamlines values of the velocity and velocity the gradients (𝑢, 𝑤, ) are obtained for the evaluation of fiber orientation. Evaluation of fiber orientation is performed outside of COMSOL using Matlab.

Evaluation of fiber orientation is performed outside of COMSOL using Matlab. Fiber orientation is evaluated using the Advani and Tucker orientation tensor evolution equation given as [2]𝑑𝑢𝑑𝑢𝑢𝑑𝑤𝑑𝑤,,,,𝑑𝑟𝑑𝑧𝑟𝑑𝑟𝑑𝑧𝐷𝑨𝐷𝑡= −12(𝜴 ∙ 𝑨 − 𝑨 ∙ 𝜴) +𝟏𝟐𝜆(𝜞 ∙ 𝑨 + 𝑨 ∙ 𝜞 − 2𝔸: 𝜞)+Dwhere the vorticity tensor is𝜴 = [(∇𝐯) − (∇𝐯)𝑇]and the strain rate tensor is𝜞 = [(∇𝐯) + (∇𝐯)𝑇]In Equation (2), A and 𝔸 are, respectively, the second and fourth-order fiber orientation tensors [2] and λ is a function of the fiber aspect ratio. The orientation diffusion term D is defined from the Folgar-Tucker Isotropic Rotary Diffusion [8] model as𝑫𝑰𝑹𝑫 = 2𝐶𝐼𝐺(𝑰 − 3𝑨)(2)(3)(4)(5)where IRD indicates isotropic rotary diffusion. In this model, 𝐶𝐼 is the empirically obtained interaction coefficient that is defined as 𝐶𝐼 = 0.075 for the present study, and 𝑰 is the identity tensor.

In this model, 𝐶𝐼 is the empirically obtained interaction coefficient that is defined as 𝐶𝐼 = 0.075 for the present study, and 𝑰 is the identity tensor. The scalar magnitude of the rate of deformation 𝐺 is defined as1 𝐺 = √ 2𝜞: 𝜞Note that Equation (2) includes the fourth-order fiber orientation tensor which is required to compute the time rate of change of the second-order orientation tensor. The fourth-order orientation tensor’s equation of motion has the sixth-order orientation tensor, and this continues for all even-ordered orientation tensors. In the present study we “close” this series using the Fast Exact Closure to approximate the fourth-order orientation tensor 𝔸.

In the present study we “close” this series using the Fast Exact Closure to approximate the fourth-order orientation tensor 𝔸. When the Fast Exact Closure [6] is used, the equations of motion presented in Equation (2) for A using Isotropic Rotatry Diffusion are rewritten as [6]𝐷𝑨𝐷𝑡= −12ℂ[𝑩 ∙ (𝜴 + 𝜆𝜞) + (−𝜴 + 𝜆𝜞) ∙ 𝑩] + 𝐷𝑟(2𝑰 − 6𝑨)𝐷𝑩𝐷𝑡= −12[𝑩 ∙ (𝜴 + 𝜆𝜞) + (−𝜴 + 𝜆𝜞) ∙ 𝑩] + 𝐷𝑟𝔻: (2𝑰 − 6𝑨)where ℂ and 𝔻 are fourth-order conversion tensors that are defined in Montgomery-Smith, et al.,[6]. The tensor B is a parameter of orientation defined similar to that of the orientation tensor A. The fiber orientation equations (7) and (8) are solved to obtain the components of A at points along the calculated streamlines. This approach provides a measure of fiber orientation at points throughout the FDM nozzle.

This approach provides a measure of fiber orientation at points throughout the FDM nozzle. In a related study we have investigated the pure shear steady state values of A and B for the given 𝐶𝐼 and 𝜆 values to provide a more accurate inlet condition and to shorten the required length of the upstream region (cf. Heller [10]). The pure shear steady state inlet condition is used for all simulations given here and the associated plots.In addition to calculating the fiber orientation in the standard nozzle appearing in Figure 1, a parametric study was performed to evaluate the effect of nozzle geometry on fiber orientation. Four parameters were selected to define different zones in the nozzle geometry as shown in Figure 3. Parameter A is the convergence zone length, B is the straight tube length, C is the amount of nozzle expansion, and D is the length of the nozzle expansion.

Parameter A is the convergence zone length, B is the straight tube length, C is the amount of nozzle expansion, and D is the length of the nozzle expansion. To assess the effect of the nozzle geometry parameters on the fiber orientation, the average A33 was computed over a cross section of the free stream flow below the nozzle exit. The x3 direction is the longitudinal flow direction of the FDM nozzle, therefore, the A33 component is a measure of fiber alignment along the axial direction of the extrudate. The average value of the fiber orientation, which is defined as 𝐴33, at steady state after the nozzle exit is defined as𝐴33 =2 (𝑟𝑜𝑢𝑡𝑒𝑟2 2 − 𝑟𝑖𝑛𝑛𝑒𝑟)𝐼where 𝑟𝑜𝑢𝑡𝑒𝑟 is the radial value of streamline 10, 𝑟𝑖𝑛𝑛𝑒𝑟 is the radial value of streamline 1, and 𝐼 is the numerically integrated area under the curve defined by the A33 values at the ten discrete streamlines in the fluid domain.

In equation (9) 𝐼 is calculated using Simpson’s 3/8 rule as(6)(7)(8)(9)𝐼 =3 8ℎ[𝐹1 + 3𝐹2 + 3𝐹3 + 2𝐹4 + 3𝐹5 + 3𝐹6 + 2𝐹7 + 3𝐹8 + 3𝐹9+𝐹10]In equation ( Fi is defined as𝐹𝑖 = 𝐴33,𝑖𝑟𝑖 𝑓𝑜𝑟 𝑖 = 1: 10and h is defined asℎ =𝑟𝑜𝑢𝑡𝑒𝑟 − 𝑟𝑖𝑛𝑛𝑒𝑟 𝑛In equation (12) n is the number of segments which is one less than the number of streamlines.The average fiber alignment, 𝐴33, is calculated for fiber orientations evaluated within each of nozzle geometry defined by the parameters appearing in Figure 3. The parametric study is performed in two steps: first, thirty-six possible individual combinations are run for parameters A and B, second, the minimum fiber alignment case for flows defined by values of A and B is then used for the computation of each of the 18 possible combinations of C and D.

From the parametric study the maximum fiber alignment case for A and B, the minimum fiber alignment case for A and B, and the max expansion case for C and D are of the greatest interest and will be discussed further in the results section.Figure 3. Nozzle Geometry Parameters and Values(10)(11)(12)ResultsFigures 4-6 below illustrate the results of our fiber orientation calculations within the nozzle flow of a typical FDM nozzle including the die swell region calculated as described above. For this study the flow domain is extended 5 exit radii past the extrusion nozzle exit to capture the effects of the extrudate swell. In Figures 4 and 5, CZS represents the start of the convergence zone of the nozzle, CZE represents the end of the convergence zone, and NE represents the extrusion nozzle exit. The location of CZS, CZE, and NE are shown in Error! Reference source not found..In Figure 4 (a), the geometry and streamlines are shown for the FDM nozzle.

The location of CZS, CZE, and NE are shown in Error! Reference source not found..In Figure 4 (a), the geometry and streamlines are shown for the FDM nozzle. Figure 4 (b) shows the A33 component for the 10 streamlines evaluated in the fluid domain. Figure 4 (b) also shows a measure of fiber alignment through the length of the nozzle and over the width of the nozzle at the 10 discrete streamlines that are plotted. This result shows the different effects that the orientation state encounters throughout the entire fluid domain.(a) (b)Figure 4. (a) Nozzle Geometry with Streamlines, (b) A33 along All Ten StreamlinesError! Reference source not found. (b) shows the effects of the FDM nozzle geometry on the fiber orientation tensor A. Error! Reference source not found. (a) shows that the A33 component of the orientation tensor increases in the convergence zone which spans from CZS to CZE for streamlines 1-8.

Error! Reference source not found. (a) shows that the A33 component of the orientation tensor increases in the convergence zone which spans from CZS to CZE for streamlines 1-8. A33 reaches a value near unity, which represents a nearly uniaxial alignment in the z-direction. The maximum value of A33 at its peak is A33 = 0.9775 along streamline 1. This increase in fiber orientation is due to elongation flow which is defined by an increase in𝜕𝑤𝜕𝑤begins to 𝜕𝑧 increase before the convergence zone which causes a similar increase in the A33 component slightly before the convergence zone. Immediately after the end of the convergence zone, in the straight portion of the nozzle from CZE to NE, there is a decrease in the A33 component which can be related to the velocity gradients shown in Error! Reference source not found. (a & d).

(a & d). The velocity gradients return to their zero in this region which causes a shear driven flow and results in the fiber orientation beginning to return to pure shear steady state orientation. After the nozzle exit, NE, there is a decrease in the A33 component of the fiber orientation which 𝜕𝑤 is due to the negative elongation component, , shown in Error! 𝜕𝑧 Reference source not found. (a & d), both of which decrease the alignment in the A33 component.as seen in Error! Reference source not found. (d). The velocity gradient,𝜕𝑧𝜕𝑤𝜕𝑢and𝜕𝑧𝜕𝑟𝜕𝑢, and an expansion flow,𝜕𝑟In Figure 4 (b) it is seen that streamlines 9 and 10 undergo decreasing alignment over a short distance in the convergence zone of the nozzle. This is due to an increase in the expansion.

This is due to an increase in the expansion . The magnitude of the flow, as seen with elongation flow quickly surpasses expansion flow and the A33 component increases similar to streamlines 1-8 until CZE.𝜕𝑢𝜕𝑤increasing faster the axial elongation gradient𝜕𝑟𝜕𝑧Figure (b) shows the A33 component of the orientation tensor throughout the fluid domain along all ten streamlines. It can be seen from Figure that the velocity gradient effects on the fiber orientation differ in magnitude depending on the spatial relation to the orientation changing flow. Because the magnitude of fluid velocity is greater at the center of the nozzle the elongation flow causes the fibers along streamline 1 to align more quickly in the axial direction than the fibers at streamline 10.

This same effect causes the decrease of fiber alignment along the axis of flow at the nozzle exit to be greater than other streamlines further from the middle of the nozzle.Figure 5 illustrates the velocity gradients along streamline 1. These velocity gradients are shown here since they directly influence changes in fiber orientation along streamline 1 throughout the nozzle. The velocities (𝑢, 𝑤) as well as the velocity gradient ( ) along streamline 1 are not included in Figure 5 as these components have been found to have very little or no effect on the fiber orientation.𝑢𝑟(a) (b)(c) (d)Figure 5. Velocity Gradients along Streamline 1Figure 6 shows values of A33 at steady state at a cross section of the flow below the nozzle exit for the ten streamlines. Values of A33 that appear in Figure 6 provide a good indication of fiber alignment in the printed strand.

Values of A33 that appear in Figure 6 provide a good indication of fiber alignment in the printed strand. With knowledge of this orientation, additional computations can be performed to determine the orientation pattern in an FDM printed part.Figure 6. A33 Components for All 10 StreamlinesIn Figure 6 it is seen that the fiber orientation is somewhat highly aligned for all values of radial position r, all showing greater than 0.55 alignment with the flow. Note that there is a high degree of alignment near the wall of the die at streamline 10 and high alignment at the center of the flow at streamline 1. There is also an obvious decline of orientation between the wall and the center of the flow. From this it can be deduced that the orientation of the fibers leaving the nozzle are highly aligned even considering the decrease in orientation right after the nozzle exit.

From this it can be deduced that the orientation of the fibers leaving the nozzle are highly aligned even considering the decrease in orientation right after the nozzle exit. This is due to the fact that the convergence zone has greatly increased the fiber alignment and that the effects of the extrudate swell cannot completely overcome the high fiber alignment.Figure 7 shows the average fiber alignment in the direction of the flow, 𝐴33, for the thirty-six combinations of the A and B geometry parameters described above. Figure 7 illustrates the effect of changing certain portions of the nozzle geometry on the average fiber alignment.

Figure 7 illustrates the effect of changing certain portions of the nozzle geometry on the average fiber alignment. From this figure the maximum fiber alignment state (A = 5, B = 0.1) and the minimum fiber alignment state (A = 5, B = 10) are determined.60.852Convergence Zone - Length 5 radiiConvergence Zone - Length 11 radiiStraight Portion of Nozzle (radii) Average A33 Value after Nozzle ExitConvergence Zone - Length 15 radiiConvergence Zone - Length 13 radii010.8Convergence Zone - Length 7 radii0.750.655Convergence Zone - Length 9 radii987100.740.93Figure 7. Average 𝐴33 Value Downstream of the Nozzle Exit for Values of A and BFigure 7 shows that parameter A, the length of the convergence zone, has very little effect on the amount of fiber alignment. The fiber alignment as measured by values of 𝐴33 varies by only 0.1 over a range of 10 radii of nozzle length B.

The fiber alignment as measured by values of 𝐴33 varies by only 0.1 over a range of 10 radii of nozzle length B. It is seen that parameter B, the length of the straight portion of the nozzle, has a much larger effect than parameter A. Parameter B decreases the fiber alignment substantially over 10 radii. This decrease in fiber alignment is due to pure shear flow in the straight portion which begins to return the orientation to the pure shear steady state orientation yielding a lower fiber alignment than that found immediately after the convergence zone ends.In Figure 8 (a), the geometry and streamlines are shown for the maximum fiber alignment FDM nozzle configuration. Figure 8 (b) shows the A33 component for all 10 streamlines in the fluid domain. Figure 8 (b) also illustrates the amount of alignment as fluid travels through the length of the nozzle. This result shows the different effects that the orientation state encounters throughout the entire fluid domain.(a) (b)Figure 8.

This result shows the different effects that the orientation state encounters throughout the entire fluid domain.(a) (b)Figure 8. (a) Maximum Fiber Alignment State Nozzle Geometry with Streamlines, (b) A33 along All Ten StreamlinesFigure 8 (b) shows that streamlines 1-5 are very similar to those in Figure 4 (b) in that fiber alignment increases slightly before CZS and then continues to increase through to CZE. Streamlines 6-10 show a decrease in alignment from shortly before CZS to slightly after CZS then increase quickly to the CZE. The larger decrease in fiber alignment and the inclusion of more streamlines is due to the lower slope of the convergence zone wall than in the previous to be greater than seen in previous nozzle calculation. This allows for the magnitude of𝜕𝑢𝜕𝑟𝜕𝑤farther into the interior of the nozzle.

This allows for the magnitude of𝜕𝑢𝜕𝑟𝜕𝑤farther into the interior of the nozzle. The straight calculations, and remain higher than portion of the tube is only 0.1 radii, therefore there is almost no decrease in fiber alignment from CZE to NE. The expansion flow , , in the extrudate swell at the nozzle exit causes a decrease in the fiber alignment as seen before, but due to the initial fiber alignment being very high the overall fiber alignment remains high after the nozzle exit.𝜕𝑧𝜕𝑢𝜕𝑤, and contraction flow,𝜕𝑟𝜕𝑧In Figure 9 (a), the geometry and streamlines are shown for the minimum fiber alignment FDM nozzle configuration. Figure 9 (b) shows the A33 component for 10 selected streamlines in the fluid domain. Figure 9 (b) also shows the amount of alignment throughout the length of the nozzle and over the width of the nozzle at the 10 discrete streamlines that are plotted.

Figure 9 (b) also shows the amount of alignment throughout the length of the nozzle and over the width of the nozzle at the 10 discrete streamlines that are plotted. This result shows the different effects that the orientation state encounters throughout the entire fluid domain.(a) (b) Figure 9. (a) Minimum Fiber Alignment State Nozzle Geometry with Streamlines, (b) A33 along All Ten StreamlinesIn Figure 9 (b) streamlines 1-10 behave similar to those in Figure 8 (b) from inlet to CZE since the two simulations have the same initial geometry. From CZE to NE there is a decrease in fiber alignment for all streamlines. This is due to pure shear flow in the straight nozzle portion which decreases the fiber alignment towards the pure shear steady state value. After NE the expansion flow and contraction flow in the extrudate swell further decreases the fiber alignment.

After NE the expansion flow and contraction flow in the extrudate swell further decreases the fiber alignment. Figure 10 shows the average fiber alignment in the direction of the flow, 𝐴33, for the eighteen possible combinations of the A, B, C, and D geometry parameters. From this figure, the maximum expansion case (A = 5, B = 10, C = 1.15, and D = 3) is selected for further study to show the effects of an expanding nozzle near its exit.1.020.591.061.140.61Expansion Length - 2 radii0.550.56Nozzle Expansion (radii) Average A33 Value after Nozzle Exit1.10.580.6Expansion Length - 1 radius1.080.571.12Expansion Length - 3 radii1.04Figure 10. Average 𝐴33 Value Downstream of the Nozzle Exit for Values of C and D, with A = 5, and B = 10Figure 10 shows that parameter C, the amount of expansion of the nozzle, decreases the fiber alignment in the nozzle.

Parameter D, the length of the nozzle expansion, is shown to have very little effect on the fiber alignment due to the tight bandwidth of the three expansion length lines. While parameter C does cause a decrease in fiber alignment it is a relatively small in comparison to parameter B. Larger values of expansion may cause a greater decrease in the fiber alignment, however, a maximum nozzle radial expansion of C = 1.15 is used here to ensure that the flow does not separate from the wall inside the nozzle.In Figure 11 (a), the geometry and streamlines are shown for the maximum expansion FDM nozzle configuration. Figure 11 (b) shows the A33 component for all 10 calculated streamlines in the fluid domain. Figure 11 (b) also shows the degree of alignment throughout the length of the nozzle and over the width of the nozzle at the 10 discrete streamlines that are plotted. This result shows the different effects that the orientation state encounters throughout the entire fluid domain.

This result shows the different effects that the orientation state encounters throughout the entire fluid domain. In this figure, STE defines the Straight Tube Exit of the nozzle. Due to the expansion zone the end of the straight tube is no longer the nozzle exit, NE.(a) (b)Figure 11. (a) Maximum Expansion Fiber Alignment State Nozzle Geometry with Streamlines, (b) A33 along All Ten StreamlinesFiber orientation along streamlines 1-10 in Figure 11 (b) show trends similar to those in Figure 9 (b) from inlet to NE due to similarities in nozzle geometry. From STE to NE a decrease in fiber alignment is seen for all streamlines. This is due to expansion flow in the nozzle expansion zone which decreases fiber alignment. After NE there is a further decrease in fiber alignment due to expansion flow and contraction flow in the extrudate swell region. This is an important finding due to the fact that most FDM nozzles are in very close proximity to the printing platform.

This is an important finding due to the fact that most FDM nozzles are in very close proximity to the printing platform. With this result the fiber orientation can be changed upstream from the nozzle exit before being deposited on the printing platform.ConclusionsIt is shown in this study that the convergence zone and extrudate swell have a large impact on the fiber orientation state for short fiber suspension in the polymer melt during FDM deposition. The effect of the extrudate swell free surface on the fiber orientation was seen to be substantial. Since the extrudate swell is the last source of major fiber orientation change in the polymer melt it has a significant effect on the final orientation of the fiber suspension which defines the material properties of the finished part.The parametric studies were run to further explore and quantify the effect of nozzle geometry on the resulting fiber alignment state.

It was seen that a change in the nozzle convergence length (Parameter A) and the nozzle expansion length (Parameter D) had very little effect on the fiber alignment in the nozzle flow. The fiber alignment was seen to undergo more substantial decreases by changing the straight tube portion (Parameter B) and the amount of nozzle expansion (Parameter C). This result allows for the augmentation of nozzle geometry to provide a preferred fiber alignment state in the printed part.This study uses the orientation tensor approach to fiber orientation in a viscous creeping flow. It is understood that the use of the Folgar-Tucker Isotropic Rotary Diffusion [5] function over-defines the rate of alignment of fibers in the transient solution, therefore in the future studies will be run using other well-known fiber orientation models to more exactly model the orientation throughout the extrusion nozzle.

Also the addition of an average fiber orientation state across the entire strand of extruded polymer melt will be of interest for future studies. The current study does offer some interesting and respectively accurate information into the effects that the convergence zone of an FDM extrusion nozzle and the extrudate swell after the nozzle exit has on the fiber orientation of chopped fibers in extruded FDM parts. The knowledge of final orientation states of fibers in the polymer melt have great significance in the calculation of mechanical properties for the final FDM printed part. Using this tool designers can perfect the printing properties for FDM printing and create optimally designed parts that are viable for industrial usesFuture WorkMany assumptions were made in this study that will affect the fiber orientation. In future studies the use of a Non-Newtonian fluid would be more accurate for the polymers that are used in FDM printing.

In future studies the use of a Non-Newtonian fluid would be more accurate for the polymers that are used in FDM printing. Also, the current research calculates the problem for an isothermal fluid flow which would need to be changed to properly define true printing conditions. Another factor that would affect the fiber orientation which needs to be considered is fully coupled fiber orientation calculation. The orientation of the fibers changes the local viscosity, and therefore would affect the velocity gradients in the fluid flow and the resulting fiber orientation state. Three dimensional finite element modelingReferences[1] Love, L. J., Kunc, V., Rios, O., Duty, C. E., Elliot, A. M., The Importance of Carbon Fiber to Polymer Additive Manufacturing, Journal of Materials Research, suppl. Focus Issue: The Materials Science of Additive Manufacturing, 29(17):1893-1898, 2014. [2] Advani, S., Tucker III, C., 1987.

1959-1970 [10] Heller, B., Effects of Nozzle Geometry and Extrudate Swell on Fiber Orientation in Fused Deposition Modeling Nozzle Flow, Master’s Thesis, Baylor University, 2015. [11] COMSOL MULTIPHYSICS, The COMSOL Group, 1 New England Executive Park, Burlington Massachusetts 018031236.

FECT OF FREE-EDGES ON MELT POOL GEOMETRY AND SOLIDIFICATION MICROSTRUCTURE IN BEAM-BASED FABRICATION OF BULKY 3-D STRUCTURESJoy E. Davis and Nathan W. Klingbeil Department of Mechanical and Materials Engineering Wright State University Dayton, OH 45435Srikanth Bontha Department of Mechanical Engineering Temple University Philadelphia, PA 19122AbstractThe success of laser and electron beam-based fabrication processes for additive manufacture and repair applications requires the ability to control melt pool geometry while still maintaining a consistent and desirable microstructure. To this end, previous work by the authors has employed point heat source solutions to investigate the effects of process variables (beam power and velocity) on melt pool geometry and solidiﬁcation microstructure (grain size and morphology) in beam- based fabrication of bulky 3-D structures. However, these results were limited to steady-state conditions away from free-edges.

However, these results were limited to steady-state conditions away from free-edges. The current work extends the approach to investigate transient behavior in the vicinity of free-edges, and follows the authors’ recent work for 2-D thin-wall geometries [1].IntroductionBoth laser and electron beam-based manufacturing are under consideration for aerospace appli- cations, for which the ability to obtain a consistent and desirable microstructure is of critical con- cern [2–6]. Speciﬁc applications for these processes include fabrication of complex parts, addition of features to existing components and repair of damage or defects. While beam-based additive manufacturing offers the promise of increased efﬁciency and ﬂexibility compared to conventional manufacturing, widespread commercialization of these processes requires the ability to predict and control melt pool size and solidiﬁcation microstructure.

In particular, the mechanical properties of the deposited material are strongly dependent on the solidiﬁcation microstructure (grain size and morphology), which is controlled by the thermal conditions at the onset of solidiﬁcation.The analysis presented herein is based on the 3-D Rosenthal solution for a moving point heat source [7]. The Rosenthal solution was ﬁrst applied to laser based deposition processes by Dykhuizen and Dobranich [8, 9]. The subsequent identiﬁcation of dimensionless process variables by Vasinonta et al. [10, 11] led to the development of thermal process maps for predicting trends in solidiﬁcation microstructure by Bontha et al. [12–16]. The Rosenthal solution is applicable for semi-inﬁnite geometries, however it does not address the response in the vicinity of a free- edge. Through direct thermal imaging of the melt pool, Rangaswamy et al.

Through direct thermal imaging of the melt pool, Rangaswamy et al. [17] observed that an increase in melt pool size occurs upon approaching a free-edge, without any change in process variables. Aggarangsi et al. [18] developed thermo-mechanical ﬁnite element models to under- stand and control the increase in melt pool size in the vicinity of a free edge. Tan et. al. [19] have recently used a superposition approach to investigate the transient temperature behavior inthe vicinity of free-edges as a function of dwell time and other process variables.

Tan et. al. [19] have recently used a superposition approach to investigate the transient temperature behavior inthe vicinity of free-edges as a function of dwell time and other process variables. It should ﬁnally be noted that nonlinear thermo-mechanical ﬁnite element modeling of laser deposition processes has also been used to predict steady-state melt pool behavior and residual stress [10, 11, 20–22], as well as to establish the utility of the Rosenthal solution for predicting trends in solidiﬁcation microstructure [12–16].In this work, the effect of free-edges on melt pool geometry and solidiﬁcation microstructure is investigated for bulky 3-D geometries, and is a direct extension of the authors’ prior work for 2-D thin-wall geometries [1]. Through superposition of the Rosenthal solution, a closed-form solution for the temperature distribution in the vicinity of a free-edge is obtained.

Through superposition of the Rosenthal solution, a closed-form solution for the temperature distribution in the vicinity of a free-edge is obtained. Melt pool contours are plotted as a function of distance from the free-edge for both small-scale and large-scale (higher power) processes. In addition, cooling rates and thermal gradients at the onset of solidiﬁcation are numerically extracted throughout the depth of the melt pool, and are subsequently plotted on solidiﬁcation maps for predicting trends in grain size and morphology for Ti-6Al-4V. Finally, ﬁnite element analysis is performed to verify the validity of the closed-form solution, as well as to investigate solidiﬁcation behavior when the beam reaches the edge and is turned off.3-D Rosenthal SolutionThis study considers the bulky 3-D geometry of Fig. 1, in which the process variables of interest are the absorbed beam power αQ and velocity V.

1, in which the process variables of interest are the absorbed beam power αQ and velocity V . It is assumed that the height h, length L and width b are sufﬁciently large such that the steady-state Rosenthal solution for a moving point heat source applies.Figure 1: Bulky 3-D GeometryFor the geometry of Fig. 1, the temperature at a point relative to the position of the point heatsource has been given in dimensionless form by Vasinonta et al. [10, 11] asT =√0+ ¯y2 ¯x20+¯z2 0)e−( ¯x0+ (cid:113) 20 + ¯y2 ¯x20 + ¯z2 0.(1)The dimensionless variables in eq. (1) are deﬁned in terms of the absorbed beam power αQ and velocity V asT =T − T0 (cid:17) (cid:16) ρcV 2k(cid:16) αQ πk(cid:17),x0 =x0 2k ρcV,y0 =y0 2k ρcVandz =z0 2k ρcV.In the above normalizations, T is the temperature at a location (x0,y0,z0) relative to the mov- ing point heat source and T0 is the initial temperature of the substrate.

The relative coordinates (x0,y0,z0) are related to ﬁxed spatial coordinates (x, y, z) at any time t as (x0,y0,z0)=(x − V t,y,z), (4), the thermophysical properties ρ, c and where V is the velocity of the beam. Also in eq. k are the density, speciﬁc heat and thermal conductivity of the material, which are assumed to be temperature-independent. In the subsequent sections, results are presented for both small-scale and large-scale processes, which are characterized in terms of the dimensionless melting temperature¯T = ¯Tm =Tm − T0 (cid:17) (cid:16) ρcV 2k(cid:16) αQ πk(cid:17).Closed-Form Solution in Vicinity of Free-EdgeFor the case of temperature-independent properties, a closed-form solution for the temperature distribution in the vicinity of the free-edge can be modeled by superposition of two point-heat sources approaching one another, with the line of symmetry representing the free-edge.

This is because the symmetry condition is equivalent to thermal insulation, which is essentially the case at the free-edge. This of course assumes that convection and radiation are negligible compared to conduction, which is generally well accepted for beam-based deposition processes.Figure 2: Free-Edge RepresentationA schematic of the superposition approach along the mid-plane y0 = 0 is shown in Fig. 2. The dimensionless distance from the free-edge is represented by the parameter a. The normalized distance a/l is used to relate the dimensionless Rosenthal melt pool length l to the distance from the free-edge, where(2)(3)l =l 2k ρcVand¯a =a 2k ρcV.It is important to note that a/l has the same value as a/l, since the nondimensionalization cancels out. As evident from Fig. 2, the effect of the second heat source can be obtained by replacing x with the quantity (2a-x) in the Rosenthal solution of eq. (1).

(6), the dimensionless variable ¯x is related to the relative coordinate x0 by the relationshipx = x0 + ¯t, where the dimensionless time ¯t is deﬁned as ¯t =t2k/ρcV 2 . The dimensionless thermalgradient is given by∇T =(cid:115)(cid:18) ∂ ¯T ∂ ¯x0(cid:19)2+(cid:18) ∂ ¯T ∂ ¯z0(cid:19)2,where ∂ ¯T ∂ ¯x0and ∂ ¯T ∂ ¯z0are obtained by differentiating eq.

Results are subsequently plotted on solid- iﬁcation maps for Ti-6Al-4V. In so doing, the thermophysical properties are assumed constant at the melting temperature Tm = 1654◦C, while the absorption coefﬁcient is assumed to be α = 0.35. Finally, the velocity is held constant at V = 8.47 mm/s. The above values are in keeping with those used in the authors’ prior work [1, 6, 12–16].Effect of Free-Edge for Small-Scale ProcessesDimensionless melt pool geometries are shown in Fig. 3 for the case of T m = 2.88. With the parameters speciﬁed above, this corresponds to a beam power of Q = 325 W. The steady-state Rosenthal melt pool is seen on the far left.

3 for the case of T m = 2.88. With the parameters speciﬁed above, this corresponds to a beam power of Q = 325 W. The steady-state Rosenthal melt pool is seen on the far left. The melt pool is an almost circular cross section and increases in size as the point heat source approaches the free-edge.Figure 3: Small-Scale Process Melt Pool Geometry, T m = 2.88(8)(9)(a) Small-Scale Process Cooling Rate(b) Small-Scale Process Thermal GradientFigure 4: Small-Scale (a) Cooling Rate and (b) Thermal GradientRepresentative results for the dimensionless cooling rate are plotted in Fig. 4 (a) as a function of relative depth within the melt pool for selected values of a/l. The relative depth within the melt pool varies in the range 0 ≤ z0/zm≤ 1, where zm is the maximum depth of the melt pool for the given value of T m. As seen in the ﬁgure, the cooling rate decreases (and ultimately becomes negative) as the point heat source approaches the edge.

As seen in the ﬁgure, the cooling rate decreases (and ultimately becomes negative) as the point heat source approaches the edge. This is in keeping with the increase in melt pool size illustrated in Fig. 3, which is somewhat analogous to an increase in incident energy. The negative cooling rate signiﬁes heating (as opposed to cooling) of the melt pool boundary, as there is nowhere for the heat to go at the free-edge. Representative results for the dimensionless thermal gradient are shown in Fig. 4 (b). The thermal gradient decreases as the beam approaches the free edge, which is again analogous to an increase in incident energy.

4 (b). The thermal gradient decreases as the beam approaches the free edge, which is again analogous to an increase in incident energy. However, the thermal gradient appears to be somewhat less sensitive to the free edge than the cooling rate, particularly near the surface of the deposit.Effect of Free-Edge for Large-Scale ProcessesFigure 5: Large-Scale Melt Pool Geometry, T m = 0.1(a) Large-Scale Process Cooling Rate(b) Large-Scale Process Thermal GradientFigure 6: Large-Scale (a) Cooling Rate and (b) Thermal GradientDimensionless melt pool geometries are shown in Fig. 5 for the case of T m = 0.1. With the parameters previously speciﬁed, this corresponds to a beam power of Q = 9332 W, which is within the range of high powers used in large-scale processes. Compared to the near-circular cross sections of the small-scale melt pools, the large-scale melt pools have an elongated, surf board- type shape.

Compared to the near-circular cross sections of the small-scale melt pools, the large-scale melt pools have an elongated, surf board- type shape. As the melt pool approaches the free-edge the size does not observably change. This is because the trailing edge of the melt pool is so far away from the free-edge that the insulated condition has little effect on the overall melt pool geometry. The dimensionless cooling rate and thermal gradient for the large-scale process can be seen in Fig. 6. As shown in the ﬁgure, both the cooling rate and thermal gradient are insensitive to the beam approaching the free-edge, which is consistent with the negligible change in melt pool geometries shown in Fig. 5.Finite Element AnalysisFinite element analysis has been used to include the nonlinear effects of temperature-dependent properties and latent heat, as well as to investigate the solidiﬁcation behavior when the beam reaches the free-edge and is subsequently turned off.

In order to reduce computation time, a 2-D axisymmetric ﬁnite element model is used instead of a 3-D ﬁnite element model. The axisymmet- ric modeling approach used herein is similar to that used for bulky 3-D geometries by Birnbaum et al. [24] and Aggarangsi et al. [25]. A representative mesh is shown in Fig. 7, and follows pro- cedures outlined in [15]. The axisymmetric axis is the axis parallel to the direction of the moving point heat source, so that the cylindrical coordinates (z, r) in Fig. 7 correspond to the Cartesian co- ordinates (x0, z0) in Fig. 1. In the axisymmetric model, the point heat source is modeled as moving through the center of a large solid that is twice the geometry actually shown. This is because the top surface of the bulky 3-D geometry is insulated, which is equivalent to a symmetry condition at z = 0. Therefore the laser power must be twice the actual for the 3-D geometry being modeled. In order to accomplish this a=0.7 is used instead of a=0.35.

Therefore the laser power must be twice the actual for the 3-D geometry being modeled. In order to accomplish this a=0.7 is used instead of a=0.35. This model allows for higher mesh resolution and shorter computation time, with no additional error.Figure 7: Representative 2-D Axisymmetric Finite Element Mesh for Transient AnalysisFigure 8: Melt Pool Contours Approaching Free-EdgeRepresentative melt pool contours are shown for the case of T m = 2.88 in Fig. 8. The results from the ﬁnite element analysis (left) compare well with results from the closed-form solution (right), which are also shown in Fig. 3.Effect of Free-Edge on Grain MorphologyAs discussed in [1, 12–16, 26], results for solidiﬁcation thermal gradient and cooling rate can be interpreted in the context of a solidiﬁcation map to provide predictions of grain size and mor- phology in Ti-6Al-4V.

Given the solidiﬁcation cooling rate ∂ T ∂t and thermal gradient G = |∇T|, the solidiﬁcation rate R is determined asR =1 G∂ T ∂t.The expected grain morphology can be predicted as either equiaxed, columnar or mixed by plotting points in G vs. R space (i.e., on the “solidiﬁcation map”), which has been previously calibrated for Ti-6Al-4V [26].(10)(a) Small-Scale Solidiﬁcation Map, T m = 2.88(b) Large-Scale Solidiﬁcation Map, T m = 0.1(c) Small-Scale FEA Results, T m = 2.88Figure 9: Solidiﬁcation Maps for (a) Small-Scale Processes (b) Large-Scale Processes and (c) Small-Scale Finite Element ResultsSolidiﬁcation maps showing the effect of free-edges on predicted grain morphology for both small-scale and large-scale deposition of bulky Ti-6Al-4V deposits are shown in Fig. 9. The values of G and R plotted in Fig. 9 (a) and (b) are extracted from the dimensionless cooling rate and thermal gradient plots of Figs.

9. The values of G and R plotted in Fig. 9 (a) and (b) are extracted from the dimensionless cooling rate and thermal gradient plots of Figs. 4 and 6, with thermophysical properties at 1654◦C and process variables as previously noted. The values of G and R plotted in Fig. 9 (c) are extracted directly from the ﬁnite element analysis, with temperature-dependent properties for Ti-6Al-4V.For the small-scale process of Fig.9 (a), the microstructure remains in the fully columnar region as the free-edge is approached. However, the free-edge acts to decrease both the solidiﬁcation rateand thermal gradient, which causes an increase in grain size. So as the free-edge is approached, results suggest a trend toward columnar grains of increasing size.For the large-scale process of Fig. 9 (b), there is very little change in solidiﬁcation microstruc- ture as the beam approaches the free-edge. This is expected based on the results of Fig. 6.

9 (b), there is very little change in solidiﬁcation microstruc- ture as the beam approaches the free-edge. This is expected based on the results of Fig. 6. For all distances from the free-edge there is a grading of the microstructure throughout the depth of the deposit, with a transition from columnar to mixed/equiaxed microstructure at the surface. This is in keeping with results of the authors’ prior work [16].From the small-scale ﬁnite element results of Fig. 9 (c), it is clear that the data points from both the steady-state and transient melt pools fall in the fully columnar region, which is in keeping with the results of Fig. 9 (a). However, results also suggest that a transition to equiaxed microstructure can occur when the beam is turned off at the free-edge (stationary melt pool). Results not shown here reveal a similar effect when the beam returns from the free-edge [23].

Results not shown here reveal a similar effect when the beam returns from the free-edge [23]. Thus, even for small- scale processes, the potential for an equiaxed microstructure exists right at the free-edge.Summary and ConclusionsThis study has considered the effect of free-edges on solidiﬁcation microstructure (grain size and morphology) in beam-based fabrication of bulky 3-D structures. Based on superposition of the steady-state Rosenthal solution for a moving point heat source, a closed-form solution has been presented for the thermal history in the vicinity of a free-edge. Dimensionless cooling rates and thermal gradients have been extracted from the closed-form solution and plotted as a function of normalized distance from the free-edge, for both small-scale and large-scale processes.

In addition, ﬁnite element analysis has been used to include the nonlinear effects of temperature-dependent properties and latent heat, as well as to investigate stationary melt pool behavior when the beam is turned off at the free-edge. Results from both the closed-form and FEA solutions have been plotted on solidiﬁcation maps for predicting grain size and morphology in Ti-6Al-4V. For small- scale processes, results suggest that melt pool geometry may be more sensitive to free-edges than the solidiﬁcation microstructure. This suggests that the control of melt pool size may be possible without affecting microstructure, which is an important result for process developers. Results further suggest that for large-scale processes, both melt pool geometry and microstructure are insensitive to the beam approaching free-edges. Finally, results suggest that a substantial free- edge effect is possible at the very end of a pass, when the beam either returns from the free-edge or is shut off.

Finally, results suggest that a substantial free- edge effect is possible at the very end of a pass, when the beam either returns from the free-edge or is shut off. In particular, results suggest that even for small-scale processes, a transition to equiaxed microstructure is possible at free-edges.AcknowledgmentsThis work has been supported by the National Science Foundation, grant numbers DMI-0224517 and CMMI-0700509, by the joint AFRL/DAGSI Research Program, project number ML-WSU- 01-11, and by Wright State University and the Ohio Board of Regents. Any opinions or ﬁndings contained herein are those of the authors, and do not necessarily reﬂect the views of the National Science Foundation, the Ohio Board of Regents or Wright State University.References[1] J. Davis, N. Klingbeil, and S.

Klingbeil, “The Role of Process Variables in Laser-Based Direct Metal SolidFreeform Fabrication,” JOM, pp. 36–39, September 2001.[6] N.W.Klingbeil, C. Brown, S. Bontha, P. Kobryn, and H. Fraser, “Prediction of Microstructure in Laser Deposition of Titanium Alloys,” in Solid Freeform Fabrication Proceedings, (Austin, TX), pp. 142–149, August 2002.[7] D. Rosenthal, “The Theory of Moving Sources of Heat and Its Applications to Metal Treat-ments,” Transactions of ASME, vol. 68, pp. 849–866, 1946.[8] R. Dykhuizen and D. Dobranich, “Cooling Rates in the LENS Process,” Sandia NationalLaboratories Internal Report, 1998.[9] R. Dykhuizen and D. Dobranich, “Analytical Thermal Models for the LENS Process,” SandiaNational Laboratories Internal Report, 1998.[10] A. Vasinonta, J. Beuth, and M. Grifﬁth, “Process Maps for Controlling Residual Stress and Melt Pool Size in Laser-Based SFF Processes,” in Solid Freeform Fabrication Proceedings, (Austin, TX), pp.

Vasinonta, J. Beuth, and M. Grifﬁth, “Process Maps for Controlling Residual Stress and Melt Pool Size in Laser-Based SFF Processes,” in Solid Freeform Fabrication Proceedings, (Austin, TX), pp. 200–208, August 2000.[11] A. Vasinonta, J. Beuth, and M. Grifﬁth, “A Process Map for Consistent Build Conditions in the Solid Freefrom Fabrication of Thin-Walled Structures,” ASME Journal of Manufacturing Science and Engineering, vol. 123, pp. 615–622, 2001.[12] S. Bontha and N. Klingbeil, “Thermal Process Maps for Controlling Microstructure in Laser- Based Solid Freeform Fabrication,” in Solid Freeform Fabrication Proceedings, (Austin, TX), pp. 219–226, August 2003.[13] N. Klingbeil, S. Bontha, C. Brown, D. Gaddam, P. Kobryn, H. Fraser, and J. Sears, “Effects of Process Variables and Size Scale on Solidiﬁcation Microstructure in Laser-Based Solid Freeform Fabrication of Ti-6Al-4V,” in Solid Freeform Fabrication Proceedings, (Austin, TX), August 2004.[14] S. Bontha, N. Klingbeil, P.

38, pp. 519–527, November 2003.[18] P. Aggarangsi, J. Beuth, and M. Grifﬁth, “Melt Pool Size and Stress Control for Laser-Based Deposition Near a Free Edge,” in Solid Freeform Fabrication Proceedings, (Austin, TX), pp. 196–207, Solid Freeform Fabrication Symposium, August 2003.[19] H. Tan, J. Chen, E. Zhang, X. Lin, and W. Huang, “Process Analysis for Laser Solid Formaing of Thin-Wall Structure,” International Journal of Machine Tools and Manufacture, vol. 50, pp. 1–8, January 2010.[20] A. Vasinonta, Process Maps for Melt Pool Size and Residual Stress in Laser-Based SolidFreeform Fabrication. PhD thesis, Carnegie Mellon University, May 2002.[21] A. Vasinonta, J. Beuth, and M. Grifﬁth, “Process Maps for Laser Deposition of Thin-Walled Structures,” in Solid Freeform Fabrication Proceedings, (Austin, TX), pp. 383–391, August 1999.[22] K. Dai and L. Shaw, “Thermal and Stress Modeling of Multi-Material Laser Processing,”Acta Materialia, vol. 49, pp. 4171–4181, 2001.[23] J.

383–391, August 1999.[22] K. Dai and L. Shaw, “Thermal and Stress Modeling of Multi-Material Laser Processing,”Acta Materialia, vol. 49, pp. 4171–4181, 2001.[23] J. E. Davis, “Effect of Free-Edges on Melt Pool Geometry and Solidiﬁcation Microstructure in Beam-Based Fabrication Methods,” Master’s thesis, Wright State University, 2010.[24] A. Birnbaum, J. Beuth, and J. Sears, “Scaling Effects in Laser-Based Additive ManufacturingProcesses,” in Solid Freeform Fabrication Proceedings, (Austin, TX), August 2004.[25] P. Aggarangsi, J. Beuth, and D. Gill, “Transient Changes in Melt Pool Size in Laser Additive Manufacturing Processes,” in Solid Freeform Fabrication Proceedings, (Austin, TX), August 2004.[26] P. Kobryn and S. Semiatin, “Microstructure and Texture Evolution during Solidiﬁcation Pro- cessing of Ti-6Al-4V,” Journal of Materials Processing Technology, vol. 135, pp. 330–339, 2003..

FECT OF GREYSCALE/PRINT DENSITY ON THE PROPERTIES OF HIGH SPEED SINTERED NYLON 12Christopher James Noble†, Adam Ellis† & Neil Hopkinson††Department of Mechanical Engineering, The University of Sheffield, Sheffield, S1 3JD, UK.AbstractHigh Speed Sintering is an Additive Manufacturing process that creates parts by sintering using inkjet and infra-red lamp technology, rather than the laser systems employed in Laser Sintering. This research investigated the effects of altering the dosage of ink (greyscale/print density) on the properties of parts produced. A clear pattern emerges that shows a ‘sweet spot’ for correct dosage of ink to maximise properties.

A clear pattern emerges that shows a ‘sweet spot’ for correct dosage of ink to maximise properties. The work also shows that greyscale allows considerable control of part density that could lead to substantial reductions in part mass beyond those that may be achieved by conventional design optimisation approaches employed today.IntroductionAdditive Manufacturing (AM) is the ‘process of joining materials to make objects from 3D model data, usually layer upon layer, as opposed to subtractive manufacturing methodologies’ [1]. It is increasingly being used for the creation of functional end-use parts, having previously been limited to producing visual representations of concepts in rapid prototyping and producing moulds in rapid tooling [2]. The layer-by-layer approach of AM combined with printing directly from computer data eliminates the need for tooling, which allows for increased design freedom and potential geometric complexity [3].

The layer-by-layer approach of AM combined with printing directly from computer data eliminates the need for tooling, which allows for increased design freedom and potential geometric complexity [3]. This also allows for optimisation and consolidation of parts, as well as being particularly relevant for individually customised parts, with even low- volume builds being economically viable.Laser Sintering (LS) is an AM process in which parts are built layer upon layer using a laser to sinter cross-sections of powdered materials. The thermal energy provided by the laser is sufficient to sinter the underlying powder. A fresh layer of powder is deposited over the sintered layer, and the process repeats until the three-dimensional part is complete. The part is then allowed to cool and post-processed, which involves cleaning the part to remove any remaining powder and any other desired finishing operations [4].

The part is then allowed to cool and post-processed, which involves cleaning the part to remove any remaining powder and any other desired finishing operations [4]. One of the biggest advantages of LS is that unlike other AM processes the powder bed is self-supporting, allowing for high levels of geometric complexity. It also gives relatively high mechanical properties, allowing its parts to be used in a wide range of applications.High Speed Sintering (HSS) is a development of Laser Sintering, but using an inkjet printhead and an IR heater rather than a laser [5]. An ink containing an infra-red radiation absorbing material (RAM) is deposited onto preheated powder via the printhead, followed by an IR lamp which passes over the entire build area. The area deposited with RAM absorbs a greater amount of thermal energy than the rest of the build bed, enough to sinter the powder. The process then repeats, layer-by-layer until the part is complete.

The process then repeats, layer-by-layer until the part is complete. HSS shares many of its advantages with LS, but eliminates some of the drawbacks. The need for expensive lasers is eliminated, machine costs are reduced and there is an increase in machine productivity as HSS cycle time isindependent of the part dimensions. The process is also scalable via the addition of printheads and IR lamps in series.The current industry standard powder is Nylon 12, a polyamide thermoplastic with a semi-crystalline structure, containing both crystalline and amorphous regions. It has been used in both LS and HSS for years, with a large processing window and relatively high mechanical properties [5].

It has been used in both LS and HSS for years, with a large processing window and relatively high mechanical properties [5]. It demonstrates a narrow melt temperature range and a low melt viscosity, meaning it can quickly reach the level of fluidity necessary for good fusion without excess energy being inputted [6].Problem DefinitionAs HSS is itself novel little research has yet been performed to investigate how print density influences the mechanical properties of parts. As the amount of energy absorbed by the powder will correspond to the density of RAM, it is expected that print density will influence mechanical properties. To date, most work has used the maximum print density the machine is capable of, the purpose of this research was to assess the influence of print density of mechanical properties of High Speed Sintered parts.Greyscale & Print DensityTo affect print density, two approaches are possible. A true greyscale image or a dithered pattern.

A true greyscale image or a dithered pattern. Greyscale offers complete coverage of the printed area with the level of greyscale is determined by the volume of the ink droplet. A dithered pattern is matrix of printed fully black dots and therefore does not cover the entire area, in this case the amount of ink deposited is determined by the spacing between dots, see Figure 1.Figure 1: Greyscale (left) vs. dithering (right)The Xaar Proton printhead used in the current HSS machine is only capable of printing in 1 bit, either a droplet is ejected or it is not. Thus, the current setup is not capable of printing variable droplet volumes. As such, a dithered pattern was used to determine print density rather than a true grayscale image.Dithering is a way of converting 8-bit greyscale images to 1-bit monochrome images, where the various levels of grey in an 8-bit image are approximated via various densities of black dots in a 1-bit image [7].

Thus, a change in greyscale level in an 8-bit image corresponds to a change in dot density in a 1-bit image. The 8-bit scale ranges from black at 0 to white at 255,therefore, the greyscale was first manipulated within this range and the subsequent grayscale image as converted into a dithered pattern as shown below in Figure 2.0285785113142170198Figure 2: The process of creating dithered patterns with variable densities, the number above represents the level of grey scale with 0 as black and 227 as a light greyThis manipulation was performed using ImageJ, an open source java based image processing software. Dithering is based on an algorithm, of which many variations exist.

Dithering is based on an algorithm, of which many variations exist. Here the chosen algorithm was Bayer 4x4, as this yielded the most even density distribution.Previous research in Laser Sintering has shown that larger particles often receive insufficient energy to fully melt, resulting in an unmolten particle core surrounded by melted and recrystallized material [8, 9]. This is evidenced by two distinct peaks in the Differential Scanning Calorimetry (DSC) curve, corresponding to melted and unmelted regions of the part, as seen in Figure 3. The amount of material which has melted and recrystallised gives the degree of particle melt (DPM).

The amount of material which has melted and recrystallised gives the degree of particle melt (DPM). As print density corresponds to the amount of RAM deposited on the surface, it was anticipated that as the print density increased this would lead to an increased amount of energy absorbed and a higher DPM.As also illustrated in Figure 3, the area under the DSC curve (mJ) may be derived, from which the heat of melting (J/g) can be calculated using the sample mass. The percentage crystallinity of the sample can then be calculated by comparing this value to that of a 100% crystalline sample, taken as 209.3 J/g based on research by Gogolewski et al [10].227Figure 3: Analysis of a DSC traceFrom this crystallinity value, the DPM can then be calculated, using Equation 1, asderived by Majewski et al.

[11].Equation 1where𝐶𝑇−𝐶𝑃 𝐶𝑀𝑅−𝐶𝑃𝐷𝑃𝑀 =is the crystallinity of the powder (0.47) andis the crystallinity of themelted and recrystallised material (0.25) and𝐶𝑃is the total crystallinity of the sample [11, 12].𝐶𝑀𝑅𝐶𝑇An increase in DPM of parts also shows a corresponding decrease in total crystallinity, due to the melted and recrystallized material having a lower crystallinity than the un-melted material (25% vs. 47%). DPM is seen to increase, and percentage crystallinity decrease, as energy density (or energy per unit area) is increased, due to a greater amount of material receiving enough energy to melt and later re-crystallise [13]. However, even when the DPM is complete, an increase in energy density continues to give an apparent decrease in percentage crystallinity [11, 12].

However, even when the DPM is complete, an increase in energy density continues to give an apparent decrease in percentage crystallinity [11, 12]. This is thought to be due to excess energy causing degradation of the polymer chains, thus destroying the long-range order of the part and reducing its percentage crystallinity [14].Previous work by Majewski et al. found that Ultimate Tensile Strength (UTS) increases with increasing energy density, due to a higher DPM, until the point where the DPM is complete, beyond which increasing energy causes a decrease in UTS values. For Elongation at Break (EaB), an increase is again seen with increasing energy density, followed by an apparent rise in EaB values at the point where the DPM is complete.

For Elongation at Break (EaB), an increase is again seen with increasing energy density, followed by an apparent rise in EaB values at the point where the DPM is complete. No clear correlation has been found between Young’s Modulus (YM) and energy density [11, 13].Experimental Methodology100% virgin Nylon 12 powder was used to manufacture 6 tensile specimens which conformed to ASTM D638-10, at the greyscale levels shown in Table 1. The machine parameters used are shown in Table 2.Greyscale Levels0285785113142170198227Table 1 - A table showing the greyscale levels usedBuild Bed (°C)Build overhead (°C)Feed bed (°C)Feed overhead (°C)Preheat stroke (% @ mm/s)Sintering stroke (% @ mm/s)171175120140100 @ 170100 @ 130Table 2: The machine parameters used to build the tensile specimensFor density and DSC analysis, a rectangular block was also built between each set of tensile specimens.

The layer thickness was 0.1 mm and the crystallinity and degree of particle melt calculated using methods described by Majewski et al.[11].Tensile tests were carried out in accordance with ASTM D638-10 standards, using a Tinius Olsen H5KS Benchtop Materials Testing Machine, extruded at a rate of 5mm/min. A stress-strain curve was obtained for each of the specimens, an example is shown in Figure 4. The mechanical properties of interest were UTS, EaB and YM.Ultimate Tensile StrengthYoung’s ModulusElongation at BreakFigure 4: An example stress strain curve produced by TiniusOlsen H5KSDensity of the samples was calculated by first obtaining the mass of each of the rectangular blocks, using an OHAUS Pioneer ™ PA64C analytical balance. A FARO Scan Arm 3D was then used to obtain a 3D scan of each of the blocks, as illustrated in Figure 5, from which part volume was obtained using Geomagic’s Studio 3D scan processing software.

A FARO Scan Arm 3D was then used to obtain a 3D scan of each of the blocks, as illustrated in Figure 5, from which part volume was obtained using Geomagic’s Studio 3D scan processing software. The density of the parts was then simply mass divided by volume.Figure 5: A typical 3D scan from the FARO Scan Arm 3DDSC measurements were recorded using a Perkin Elmer DSC 8500 machine, and analysed using Pyris DSC analysis software. In accordance with previous work by Zarringhalam a 12mg sample (±3mg) was heated from 20oC to 220 oC at a rate of 10 oC/min, before being held at 220 oC for two minutes, and then being cooled back down to 20oC at the same rate of 10 oC/min. The samples were cut from the centre of each of the rectangular blocks [12].Results and DiscussionFigure 6 shows how ultimate tensile strength is influenced by greyscale.

The samples were cut from the centre of each of the rectangular blocks [12].Results and DiscussionFigure 6 shows how ultimate tensile strength is influenced by greyscale. It is important tonote that a lower greyscale value indicates a higher print density.5045403530252015105250200Greyscale Level 1501005000Figure 6: UTS vs. Greyscale Level) a P M( S T UAn increase in UTS is observed with a decrease in greyscale until the 142 greyscale level point. Beyond this, a decrease in UTS with decreasing greyscale was observed. The initial increase is likely to be due to a more complete fusion and better particle-to-particle contact, due to the increased energy absorbed due to more RAM.

The initial increase is likely to be due to a more complete fusion and better particle-to-particle contact, due to the increased energy absorbed due to more RAM. The decrease after the 142 greyscale level can be attributed to degradation of the polymer material and a likely decrease in part crystallinity, with lower crystallinity values associated with lower UTS values [15].30252015105250200Greyscale Level1501005000Figure 7: EaB vs. Greyscale LevelFor EaB, as seen in Figure 7, a clear initial increase is observed once again, likely due to better fusion and particle-to-particle contact. There then appears to be a rise in EaB values around the 142 greyscale level, similar to results observed by Majewski et al.

There then appears to be a rise in EaB values around the 142 greyscale level, similar to results observed by Majewski et al. Beyond the 142 greyscale levels, a clear trend becomes harder to discern; it appears the results are broadly flat, with natural variations in the data.)%( B a E18001600) a P M140012001000800600(s u l u d o M s ' g n u o Y400200250200Greyscale Level 1501005000Figure 8: YM vs. Greyscale LevelFor YM, an initial increase with decreasing greyscale can be observed in Figure 8, likely due to similar reasons as described previously. Beyond this however there is no clear trend, as expected based on previous research.12001000800600) 3m / g k ( y t i s n e D400200250200Greyscale Level 1501005000Figure 9: Density vs. Greyscale LevelIn terms of part density, as seen in Figure 9, a sharp initial increase is seen with decreasing greyscale level, until the 170 greyscale level.

Greyscale LevelIn terms of part density, as seen in Figure 9, a sharp initial increase is seen with decreasing greyscale level, until the 170 greyscale level. Beyond this, the increase in density with decreasing greyscale is not so pronounced. This is likely due to the DPM being almost complete at this point, with any further energy just helping to close up any remaining pores. Combinedwith the previous results, a clear potential can be seen to alter density of components whilst retaining acceptable mechanical properties.Figure 10 shows the percentage crystallinity of each set of samples. As expected, a decrease in crystallinity with decreasing greyscale is seen, initially due to the greater amount of melted and recrystallized material, then due to degradation of the polymer chains destroying the long range order of the parts.3530252015105250200Greyscale Level1501005000Figure 10: Crystallinity vs.

Greyscale LevelFrom the crystallinity trendline, the DPM could be derived at each of the greyscale levels, giving the results seen in Table 3. The DPM appears to be complete somewhere between the 170 and 142 greyscale levels, which corresponds well with results seen previously; the highest UTS value was observed around the 142 greyscale level, and a jump in EaB occurred between 170 and 142. This is similar to the results observed by Majewski et al.Greyscale Level 227 198 170 142 113 85 57 28 0Derived DPM (%) 92.01 95.14 98.15 101.17 104.30 107.31 110.33 113.44 116.47Table 3: The change in greyscale level vs. derived DPM)%( y t i n i l l a t s y r COne interesting point to note is that the DPM continues to rise above 100%, which is clearly illogical. This is due to a fundamental limitation in Equation 2, whereby it is only valid for crystallinity values above the crystallinity of the melted and recrystallized region (25%).

This is due to a fundamental limitation in Equation 2, whereby it is only valid for crystallinity values above the crystallinity of the melted and recrystallized region (25%). In practice, crystallinity can fall below this due to degradation of the polymer material. Therefore, Equation 1 should be considered an identifier of where the DPM is complete, rather than as an exact calculation tool.ConclusionsThe experiments have shown that the amount of RAM has a significant effect on the properties of High Speed Sintered parts, with a decrease in greyscale/print density increasing the amount of thermal energy absorbed. The degree of particle melt at these parameters is complete around the 142 greyscale level, where maximum Ultimate Tensile Strength values were observed and very high relative values of Elongation at Break and Young’s Modulus.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of Heat Treatment and Hot Isostatic Pressing on the Morphology and Size of Pores in Additive Manufactured Ti-6Al-4V PartsBin Zhang1, Kyungmin Ham2, Shuai Shao1*, Nima Shamsaei3*, Scott M. Thompson31 - Department of Mechanical and Industrial Engineering, Louisiana State University, Baton Rouge, LA 70803 2 - Center for Advanced Microstructures and Devices, Louisiana State University, Baton Rouge, LA 70803 3 – Laboratory for Fatigue & Additive Manufacturing Excellence, Department of Mechanical Engineering, Auburn University, Auburn, AL 36849 * Corresponding Authors: sshao@lsu.edu, shamsaei@auburn.edu,AbstractAdditive manufactured parts suffer from porosity, among other defects. The slit-shaped pores due to lack of fusion are the most detrimental to fatigue and mechanical properties.

The slit-shaped pores due to lack of fusion are the most detrimental to fatigue and mechanical properties. Their sharp edges generate severe stress concentration and serve as preferred sites for crack initiation. The sharp edges also have low formation energy of surface adatoms, increasing their tendency to spheroidize under elevated temperatures, such as during heat treatment (HT). In hot isostatic pressing (HIP), the combined action of high temperature/pressure also reduces the size of the pores. In this work, HT and HIP were performed on Ti-6Al-4V parts manufactured from laser- based powder bed fusion to investigate the effect of HT and HIP on morphology/size of pores. Using scanning electron microscopy combined with X-ray computed tomography, special attention is paid to the evolution of the shape of the poros under controlled exposures to elevated temperature during HT.

The results will be used, in our subsequent work, to validate a phase field porosity evolution model based on density functional theory.IntroductionAdditive manufacturing (AM) is extremely prone to the formation of pores in parts, which include the general spherical pores as well as the slit-shaped, lack-of-fusion pores (LoFP). The latter, generally resulted from insufficient energy input, represent perhaps the most detrimental defects to the mechanical properties as well as fatigue properties of AM parts. The LoFP have sharp edges which correspond to very high stress intensity factors which can significantly reduce the ductility of the AM parts when the LoFP are large [1]. Even when the LoFP are small, the high stress concentration ahead of the tip of a LoFP can lead to accelerated initiation and propagation of fatigue cracks under cyclic loading.

Even when the LoFP are small, the high stress concentration ahead of the tip of a LoFP can lead to accelerated initiation and propagation of fatigue cracks under cyclic loading. As a result, the fatigue life of an AM part can be reduced by as much as a factor of ten compared to conventional wrought parts [2].Nowadays, although the process parameters of state-of-art additive machines can be fine- tuned for relative density of well above 99.5% [3], the porosity cannot be completely eliminated. However, even at such near full-dense conditions, the shape of the pores cannot be regulated and the appearance of LoFP persists [4]. Hot isostatic pressing (HIP) has been shown as a very promising process to remove the LoFP, or porosity general, from the AM parts. HIP condenses parts by exposing parts in a high temperature and high pressure Ar environment, and has been107proven effective in closing pores in parts produced from electron beam type of AM processes.

In laser-based AM parts, due to the employment of non-soluble gases (such as Ar) as delivery/shielding gas, the gas entrapped pores are not easily removed by HIP [5,6] and lead to inconsistent fatigue performances [2].Due to the increased mass diffusion coefficient, actelevated temperature alone can cause the LoFP to change shape, even without the assist from the hydrostatic pressure. The sharp edges of the LoFP have low formation energy of adatoms on surface as well as vacancies in bulk and therefore act as strong sink for surface/interstitial atoms and active source for vacancies. As a result, they have strong tendency to spheroidize under elevated temperatures. Driven by the same mechanism, cracks in ceramics can be healed when annealed for an extended period [7]. Therefore, the sharp edges of LoFP are expected to be “rounded” when the pore bearing AM part is heat treated.

Therefore, the sharp edges of LoFP are expected to be “rounded” when the pore bearing AM part is heat treated. The resulting part, although none of its pores are expected to close, might possess improved mechanical/fatigue performance due to the significantly alleviated stress concentration ahead of the LoFP. However, to the authors’ best knowledge, the effect of HT on the shape of the LoFP, or as we put it: the pore conditioning effect, in AM parts have not been addressed before.In this work, we attempt to investigate the effect of HT and HIP on the size and morphology of pores. Small cylindrical Ti-6Al-4V (Ti-64) samples are manufactured from laser based powder bed fusion AM machines and are subjected to HIP and HT. After the treatments the porosity of the samples are evaluated using X-ray computed tomography (XCT) and scanning electron microscopy. The ability of HIP in terms of pore closure as well as the ability of HT in terms of pore conditioning are examined.

The ability of HIP in terms of pore closure as well as the ability of HT in terms of pore conditioning are examined. Specially, we pay specially attention to the effect of elevated temperature on the change of sharpness, measured using the radius of curvature, ahead of the LoFP. The results obtained from this work, combined with our previous result from density functional theory calculations (DFT), will be used to inform/validate a phase field model to predict the evolution of porosity under HT as well as HIP.Experimental SetupA total of seven nominally identical, rod samples are manufactured using the EOS M290, for which the Ti64_030_100 PerformanceM291 material parameter set was chosen in order to minimize porosity. The default InnerCore parameters were used, which have a laser power of 170 W, a travel speed of 1250 mm/s, a hatch spacing of 0.1 mm, and a layer thickness of 30 µm.

The default InnerCore parameters were used, which have a laser power of 170 W, a travel speed of 1250 mm/s, a hatch spacing of 0.1 mm, and a layer thickness of 30 µm. After each layer, two post contour exposures took place: one at 150 W, 1250 mm/s, and a 0.020 mm offset and a second one with the same laser power and speed, but with a 0 mm offset. These post contour exposures only expose the outside edges of the specimen, and serve to smooth any edges or remove un-melted powder attached to the outer surface. After the post contour, the build platform is moved down, a new layer of powder is dispensed, and the next layer is fabricated. The hatching orientation of each layer is rotated by approximately 67° with respect to the deposition orientation of the previous layer. The samples were supported with 3mm default block supports, with fragmentation disabled, in order to facilitate easy removal from the build substrate.

The samples were supported with 3mm default block supports, with fragmentation disabled, in order to facilitate easy removal from the build substrate. During the builds, Ar is used as shielding gas to prevent oxidization.One sample is sent to Quintus Technologies, LLC [8] for HIP treatment. The treatment consists of a ramp-up stage of both temperature and pressure to working conditions, i.e. 820°C and 200 MPa, a holding time of 2 hrs, rapid cooling, followed by depressurization. The working gas is108Ar. Simple heat treatments with near-constant temperatures (950°C) and various treating times (3hr, 6hr, 12hr, 24hr, 48hr) are performed on five samples. The temperature is set below the β transus temperature (~1000°C) [9], so that the β grain structure is not destroyed by HT. All samples are individually wrapped using stainless steel foils heat treated in the same furnace. At each of the aforementioned time, a sample is taken out of the furnace.

All samples are individually wrapped using stainless steel foils heat treated in the same furnace. At each of the aforementioned time, a sample is taken out of the furnace. Surface oxidization of different degrees are found on the samples.Figure 1. Samples mounted on epoxy, ground and polished. The samples and their specific treatments have been designated in the figure. 0h, 3h … denote the length of HT time. Black arrows show the build direction.The samples, before and after HT, are scanned by XCT in the tomography beamline of the synchrotron located in Center for Advanced Microstructures and Devices (CAMD), Louisiana State University. Scintillators used are 500µm thick YAG(Ce) and 250µm thick LuAG, yielding X-ray energy of 38keV. Camera used are PCO.edge 5.5/4.2 with 6.5µm pixel size, the resulting effective voxel size is ~4 µm. X-ray images are taken at angle increment of 0.5°, in a 180° range. At each angle, four images are taken, each with 2 s exposure time.

X-ray images are taken at angle increment of 0.5°, in a 180° range. At each angle, four images are taken, each with 2 s exposure time. After the XCT scans, sample are then mounted in epoxy, longitudinally ground to the midsection with sand paper, then polished to a mirror finish with diamond solution. The morphology of the pores was examined by Scanning Electron Microscopy (SEM) at 30 KV.Results and Discussiona. Evaluation of Porosity on As-built SamplesTo get an initial assessment of porosity level in the AM Ti64 parts, XCT scans are performed on all seven samples. During the XCT scans, the samples are mounted “upside-down” where the top portion of the samples are gripped. A typical XCT scan result is presented in Fig. 2, which shows the 3D reconstructed data of Sample c. In attempt to reveal the internal defect structure of the samples, the bulk of the sample is removed and only surfaces are displayed. In Fig.

2, which shows the 3D reconstructed data of Sample c. In attempt to reveal the internal defect structure of the samples, the bulk of the sample is removed and only surfaces are displayed. In Fig. 2(a), the dashed ellipse marks the bottom surface of the sample, which contains significant roughness as is evident from Fig. 1(c). The circumferential surface of the sample display109significantly less roughness thanks to the post contour exposures during AM. Fig. 2(b) shows the internal of the sample through the bottom view of the 3D reconstructed data. The concentric rings are artifacts created during the XCT scan, possibly due to the dust collected on the scintillator as well as the COMS sensor. Other resolved “defects” (indicated by white arrows) are from the bottom surface of the sample. It is evident from Fig. 2(b) that the XCT revealed no apparent porosity on Sample c. Porosity was also not detected by XCT in the rest of the samples, both before and after the HT.Figure 2.

It is evident from Fig. 2(b) that the XCT revealed no apparent porosity on Sample c. Porosity was also not detected by XCT in the rest of the samples, both before and after the HT.Figure 2. A typical 3D reconstruction of a XCT scan on Sample c before HT.To further evaluate the porosity character in the as-built conditions, Sample b is longitudinally cut/polished and examined under SEM (Fig. 3(a)-(c)). As shown, both LoFP and spherical pores with microscopic sizes present. The thickness of the defects is only ~5 µm which cannot be resolved by XCT due to the back ground noise. The sharpest edge of the LoFP shown in Fig. 3a has a radius of curvature of only 75 nm. Based on the area fraction of pores on the cross sectional SEM images, the part density of all the samples (including Sample (c)) is estimated to be >99.9%.Figure 3.

Based on the area fraction of pores on the cross sectional SEM images, the part density of all the samples (including Sample (c)) is estimated to be >99.9%.Figure 3. SEM images showing cross section of pores in (a-c) as-built Sample c and (d) Sample d after 3h HT.110It is also interesting to note that there a blemish near the top of the Sample c as shown in Fig. 1(c) (indicated by red arrow). However, this portion of the sample is gripped while the X-ray images are taken, therefore this pore is not revealed in the XCT scans. SEM confirms this pore as a large LoFP (over 1 mm in size) with un-melted powder (Fig. 3(d)).b. Effect of HT on the Shape of PoresUnder elevated temperatures, the sharp edges of the LoFP show increased tendency of spherodization as a result of the increased mass diffusion coefficient. The XCT, limited by its resolution, is insufficient to characterize this phenomenon. Instead, we use a non-exhaustive approach based on SEM imaging.

The XCT, limited by its resolution, is insufficient to characterize this phenomenon. Instead, we use a non-exhaustive approach based on SEM imaging. For each sample shown in Fig. 1, two longitudinal cross sections near the midsection are polished and imaged in attempt to locate the LoFP and their sharp edges. The smaller principal radius of the sharp edges of the LoFP are estimated based on the cross section of the pores on the exposed cross sections. This methodology, while not capable of capturing the evolution of pores in a given sample, does provide an upper bound evaluation on the minimum radius of curvature (MRC) for all LoFP within a sample. Fig. 4 shows the MRC in the as-built sample (Fig. 1(b)) as well as the samples heat treated for 3h, 12h and 24h (Fig. (c), (e) and (f)). A clear increasing trend of the MRC can be observed when the HT time is below 24h – an increase from a mere 75 nm to almost 2 μm (25 fold increase).

(c), (e) and (f)). A clear increasing trend of the MRC can be observed when the HT time is below 24h – an increase from a mere 75 nm to almost 2 μm (25 fold increase). Assuming constant LoFP length, this increase in the MRC lead to a reduction of stress concentration factor of 5 [10]. However, with prolonged HT of 48h, the MRC reduces to a lower value, associated with the formation of facets at the tip of LoFP (dashed lines in Fig. 4).Figure 4. Minimum radius of curvature of LoFP as a function of HT time. Insets show the morphology of the sharpest tips of LoFP in respective samples in the same magnification.Under HT for 48h, the formation of the facets also manifests in spherical pores and un- melted powder particles, such as ones shown in Fig. 5. Note that although the overall dimension of the LoFP shown in Fig. 5(b) is of tens of μm, its width is only 10 μm or below. As such, the XCT is not able to resolve this type of defect.

5. Note that although the overall dimension of the LoFP shown in Fig. 5(b) is of tens of μm, its width is only 10 μm or below. As such, the XCT is not able to resolve this type of defect. Faceting is thermodynamically driven as the α- and β- Ti has preferred, low energy surfaces. For instance, density functional theory calculations and {0001} surfaces of α Ti have free energies of 1.874 and 1.938 [11,12] reveal that the J/m2 and are preferred. The {110} surfaces of β Ti has free energy of only 0.68 J/m2 and are preferred. This trend is expected to be valid, although the presence the alloying elements of Al and{112�0}111V may slightly affect the numerical values of the surface energies. According to the free surface energies, the equilibrium crystal shapes of the α- and β- Ti are obtained using the well-known Wulff construction [13] and shown in Figs. 5(c) and (d).Results presented in Figs.

5(c) and (d).Results presented in Figs. 4 and 5 suggest that, although HT may be able to improve the mechanical/fatigue performance of the AM parts by blunting the edges of LoFP, the extent of improvement offered by HT does not monotonically increase with increasing HT time. Instead, an optimum window exits, beyond which the edges of the LoFP deteriorate with concomitant formation of facets.Figure 5. Faceting of pores (a) as well as un-melted powder particles (b) under HT for 48h. The Wulff construction of equilibrium crystal shape of the α- and β- Ti are shown in (c) and (d).c. Effect of HIP on PoresSEM examination on Sample (a) did not reveal any apparent porosity. Our previous theoretical model has reveal that, under HIP conditions, the size of pores containing Ar gas are expected to reduce by two orders of magnitude [6]. A pore with diameter of 50 μm can reduce to a diameter of only 500 nm.

A pore with diameter of 50 μm can reduce to a diameter of only 500 nm. This makes the probability of exposing the pores by means of polishing extremely low. In our future work, the build parameters will be purposefully adjusted so that large pores (with diameters of around 1mm) are present in samples leading to easily detectable porosity even after HIP. In doing so, the effect of HIP on both the size as well as the shape of the pores can be evaluated in detail.ConclusionIn this work, using scanning electron microscopy combing X-ray computed CT, we attempt to investigate the effect of heat treatment and hot isostatic pressing on the size and morphology of pores in parts manufactured from laser powder bed fusion. It was shown that the EOS M290 additive machine with the Ti64_030_100 PerformanceM291 material parameter set is able to produce parts with density well above 99.9%. However, the large fraction of lack-of-fusion pores112exist.

However, the large fraction of lack-of-fusion pores112exist. Even though the length of the typical lack-of-fusion pores can be near 100 μm, their thickness is usually one order of magnitude smaller. Therefore the porosities could not be resolved by the XCT scans performed at CAMD.As revealed by SEM performed on the longitudinal cross sections of the samples, heat treatments has an interesting blunting effect on the sharp edges of lack-of-fusion pores. This effect may significantly improve the mechanical/fatigue performance of AM parts, as the stress concentration at the edges are alleviated. However, prolonged heat treatment may also re-sharpen the blunted edges associated with the formation of low energy surface facets.

However, prolonged heat treatment may also re-sharpen the blunted edges associated with the formation of low energy surface facets. SEM did not reveal any porosity on the sample treated by hot isostatic pressing, which may be ascribed to the sub- micron pores that are difficult to be exposed by polishing.The findings obtained in this work suggest that, in order for models to effectively predict the evolution of porosity in AM parts under heat treatment and the resulting mechanical/fatigue performance, it is necessary to incorporate the effect of the anisotropic free surface energies.AcknowledgementS. Shao acknowledges the support provided by the Office of Research and Economic Development and the Department of Mechanical and Industrial Engineering, Louisiana State University. Partial funding from the Naval Air Systems Command (NAVAIR) and Magee Technologies LLC is greatly appreciated.References[1] B. Torries, S. Shao, N. Shamsaei, S.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEFFECT OF HEAT TREATMENT OF 316L STAINLESS STEEL PRODUCED BYSELECTIVE LASER MELTING (SLM)M. L. Montero Sistiaga*, S. Nardone †, C. Hautfenne† andJ. Van Humbeeck *KU Leuven, Department of Materials Engineering, Heverlee B-3001, Belgium. †Engie Lab, Linkebeek B-1630, BelgiumAbstractSelective Laser Melting (SLM) shows a big potential within additive manufacturing of metals. The competitive mechanical properties compared to conventional processes as well as the geometry freedom are the main advantages of SLM. 316L stainless steel has been investigated in previous works regarding microstructure and mechanical properties. However, the influence of heat treatments has not been fully reported yet.

316L stainless steel has been investigated in previous works regarding microstructure and mechanical properties. However, the influence of heat treatments has not been fully reported yet. This work studies the influence of different heat treatments applied to 316L stainless steel produced by SLM. The microstructure evolution was investigated for different conditions. Tensile, Charpy and hardness tests were performed on the as built and heat treated samples.Keywords: Selective Laser Melting, Additive Manufacturing, 316L, heat treatmentIntroductionSelective Laser Melting (SLM) is an Additive Manufacturing process which locally melts a metallic powder bed using a highly focus laser beam. The high cooling rates result in a unique microstructure, but may introduce residual stresses making the material susceptible to distortion [1].

The high cooling rates result in a unique microstructure, but may introduce residual stresses making the material susceptible to distortion [1]. The high flexibility in design, low material waste and fast production of near-net-shape parts are the main advantages compared to conventional processing routes.316L stainless steel has been widely investigated for SLM. High relative densities and competing mechanical properties have been achieved. The optimization process for 316L has been reported by many authors[2], [3]. The work of Yasa et al. showed that re-melting results in an increase of density and surface quality. After SLM a sub-micron cellular-dendritic solidification microstructure is observed caused by the high thermal gradients during cooling. These cells are oriented more or less towards the heat gradient inside the melt pool[4], [5].

These cells are oriented more or less towards the heat gradient inside the melt pool[4], [5]. Referring to the mechanical properties, the yield strength of SLM processed 316L parts appear to be much higher than wrought 316L, and a large difference in elongation is found between horizontal and vertical built samples [6], [7]. Due to the presence of large thermal gradients and the local melting and solidification processes during SLM, microstructural and mechanical558anisotropy are inherent to happen [6], [7]. The fatigue properties of SLM processed 316L parts have been investigated[8], [9].This works aims at better understanding the effect of the heat treatments on the microstructural and mechanical behavior of 316L parts produced by SLM.

Tensile and Charpy test were performed to investigate the evolution in mechanical properties after applying different heat treatments.Materials and methodsThe material used in this study was 316L stainless steel provided by SLM solutions GmbH and with powder particles sizes ranging from 10 to 45 µm. The 316L powder composition is shown in Table 1as defined in ASTM B243. The powder morphology is shown in Figure 1. The particles show mainly spherical shaped granules with some smaller satellite particles attached.Table 1. Chemical composition in weight % of 316L powder.Weight [%] 316L powderFeBal.Cr16.8Ni10.4Mo2.1Mn1.11Si0.56P0.03SC0.011 0.01Density [g/cm3]7.985The SLM 280 selective laser melting machine was used for the production of the 316L samples. The SLM instrument is equipped with 400W fiber laser.All the parts were built under argon atmosphere to avoid oxidation and using a pre-heating temperature of 100°C.

The SLM instrument is equipped with 400W fiber laser.All the parts were built under argon atmosphere to avoid oxidation and using a pre-heating temperature of 100°C. A relative density of around 99.3% was achieved for all samples measured by Archimedes method using a theoretical density of 7.985 g cm-3.For microstructural observations the samples were grinded, polished and chemically etchedduring 30s using 10% oxalic acid dissolved in H2O.The samples were examined using an Axioskop 40Pol/40 A Pol optical microscope (LOM) and Philips XL30FEG Scanning ElectronMicroscope (SEM). The crystallographic orientation was studied using electron backscattered diffraction (EBSD) using aPhilips XL30 SEM equipped with a TSL orientation imaging microscope system.559Figure 1. 316L stainless steel powder morphology (a). Part positioning for the tensile (b) and Charpy (c) test.Rectangular tensile test specimens were built in two directions XY and XZ (see Figure 1b).

316L stainless steel powder morphology (a). Part positioning for the tensile (b) and Charpy (c) test.Rectangular tensile test specimens were built in two directions XY and XZ (see Figure 1b). The samples were tested in an Instron 4505 at a cross head velocity of 1 mm min-1 according to ASTM E 8M.Rectangular bars with size 12x12x56 mm3 were horizontally built and later machined for the Charpy V-notch test according to ASTM E23. Two different directions were tested with the notch in XY and XZ direction (see Figure 1c).The Vickers hardness tester FV-700 was used for measuring the Vickers hardness with 0.5 kg load for 15s with five repetitions per sample.Heat treatments were performed using a vertical tube furnace under argon atmosphere and with a heating rate of 10°C min-1. The different treatments are depicted in Table 2. HT1 and HT2 apply air cooling and HT3 applies water quenching.Table 2.

The different treatments are depicted in Table 2. HT1 and HT2 apply air cooling and HT3 applies water quenching.Table 2. Heat treatments applied to the 316L tensile and Charpy samples.Heat treated (HT) HT1 HT2 HT3HT cycle 600°C 2h, air cooling 950°C 2h, air cooling 1095°C 2h, water coolingResults and discussion1. Microstructure evolution after heat treatmentsFigure 2shows the side view microstructures for the as built (AB) and heat treated samples by SEM. For the AB and HT1(Figure 2a-b) condition, melt pool boundariesare visible. The grains grow across these melt pools parallel to the building directions, in the direction of the heat gradient. Inside the grains, a clear cellular dendritic microstructure is observed, with cell sizes ranging between 0.5 and 1 µm. The microstructures observed by SEM for samples HT2560(Figure 2c) and HT3 (Figure2d) present no cellular dendrites inside the grains.

The microstructures observed by SEM for samples HT2560(Figure 2c) and HT3 (Figure2d) present no cellular dendrites inside the grains. The melt pool boundaries are almost dissolved and only the grains are observed. After etching, some homogeneously distributed round pits appear inside the grains for HT2 and HT3 at higher magnifications. Although no substructure can be observed for HT2 and HT3 (e.g. cellular dendrites) small chemical variations within the grain could act as initiation sites of pitting. The difference in pit size between HT2 and HT3 can be related to the etching time varianceFigure 2 Secondary electron image of the side view of 316L stainless steel produced by SLM. (a) As built (AB), (b) HT1, (c) HT2 and (d) HT3 conditions. The arrows are indicating the melt pool border, present for AB and HT1 samples.Figure 3.

(a) As built (AB), (b) HT1, (c) HT2 and (d) HT3 conditions. The arrows are indicating the melt pool border, present for AB and HT1 samples.Figure 3. EBSD orientation maps of samples AB (a), HT1 (b), HT2 (c) and HT3 (d), plane parallel to the building direction.Figure 3 depicts the EBSD image of the side view for the AB sample. The results are in consistence with the SEM observations. Columnar grains grow parallel to the building direction561with longitudinal size around 100 µm. The EBSD results conducted for HT1, HT2 and HT3 shows no difference in grain size in comparison with the AB samples. Thus the heat treatments applied to the as built samples do not show grain enlargement. All the samples show grains directed parallel to the building.2. Mechanical PropertiesVickers micro-hardness measurements are presented in Table 3. The hardness for conventionally produced 316L is also listed as a reference.

Mechanical PropertiesVickers micro-hardness measurements are presented in Table 3. The hardness for conventionally produced 316L is also listed as a reference. The experimental results obtained for the SLM processed parts in AB and heat treated condition are well above the requirements. For the AB condition, a hardness of 245 ± 25 HV was measured. The AB hardness is in accordance with the values found in literature, with hardness values around 235HV [4], [6]. HT1, which applies only a heat treatment at 600°C, increases the AB hardness up to 271 ± 25 HV. HT2 and HT3 creates lower values than AB and HT1.Table 3. Vickers micro hardness measurements of the side view of the SLM processed samples.Condition Vickers hardness [HV]AB 245±21HT1 271±25HT2 215±14HT3 212±20Ref. [10] 155The tensile test and Charpy-V test were performed on the as built and heat treated samples.

[10] 155The tensile test and Charpy-V test were performed on the as built and heat treated samples. The results are shown in Figure 4.EM10216-5:2013 standard is also listed as reference for the elongation, yield tensile strength (YTS) and ultimate tensile strength (UTS) values. All samples exhibit a much higher YTS compared to the standard. For the UTS all values lie within the limitations of the standard. On the other hand, varying elongation values with wide standard deviations are observed. The porosities or impurities found after the SLM process can deteriorate the ductility values, but they do not affect other properties such as UTS. Nevertheless, most of the samples lie close to the EN 10216-5:2013 standard.AB and HT1 samples exhibit a higher yield and ultimate strength compared to HT2 and HT3. According to the grain boundary strengthening the grain size influences the yield strength in 316L[4].

According to the grain boundary strengthening the grain size influences the yield strength in 316L[4]. HT2 and HT3 present a grain size similar to the AB and HT1, but no cellular dendritic substructure (see Figure 2).This suggests that the high strength obtained by AB and HT1 is attributed to the presence of a cellular dendritic substructure. The dislocations could accumulate at boundaries of the cells, acting as a strengthening sites [4]. This also can explain the high hardness (see Table 3) obtained for samples AB and HT1 compared to HT2 and HT3.562Figure 4 Mechanical properties of 316L in as built (AB) and heat treated conditions (HT1, HT2, HT3) and compared with EN 10216-5:2013 standard.Charpy test results are presented in Figure 4d for the AB and HT1-3 samples. HT2 shows the highest absorbed energy with 154 ± 6 J for XY direction and 152 ± 8 J for XZ direction.

HT2 shows the highest absorbed energy with 154 ± 6 J for XY direction and 152 ± 8 J for XZ direction. HT1 and HT3 samples show similar toughness compared to the AB condition.For HT2 the reduction of tensile strength maintaining the elongations observed in the tensile values results in an improved toughness compared to the AB condition. On the other hand, HT2 and HT3 show similar microstructure as seen in Figure 2, but significantly different absorbed energies. HT3 is subjected to a dwell temperature higher than HT2 during the same time (2h) and the sampleis subsequently water quenched. The reduced toughness obtained for HT3 could be accounted forthe effect of water cooling, which could induce stresses to the material.

The reduced toughness obtained for HT3 could be accounted forthe effect of water cooling, which could induce stresses to the material. However, the influence of air cooling and water quenching after the same heat treatment should be further investigated.All the values obtained by the Charpy test (absorbed energy) obtained in this work show higher values than the ones found in literature for SLM of 316L.563The fracture surfaces of the Charpy samples were analyzedin order to understand the differences in mechanical properties after the different heat treatments. Figure 5 displays the fracture surfaces for all the samples. In Figure 5e, the very ductile fracture can be observed, visible for all the condition. At a closer look with the SEM, the classic dimple structure can be observed. No non-molten particles can be observed for all samples, proving the good consolidation of the parts.

At a closer look with the SEM, the classic dimple structure can be observed. No non-molten particles can be observed for all samples, proving the good consolidation of the parts. Sample HT3 (Figure 5d) shows wider and shallower dimple structures, which could explain the lower impact toughness compared to HT2. For samples AB, HT1 and HT3, the dimple sizes are comparable to the one of the cells. The observed dimples show a size between 0.5 and 1 µm, even for HT3 where no cellular dendrites where observed.Figure 5. Fracture surfaces of SLM processed 316L after Charpy test using scanning electron microscope for (a) as built, (b) HT1, (c) HT2 and (d) HT3 samples.

Fracture surfaces of SLM processed 316L after Charpy test using scanning electron microscope for (a) as built, (b) HT1, (c) HT2 and (d) HT3 samples. Optical image of side and front view of the surface fracture, showing a ductile mode (e).Conclusions316L stainless steel processed by SLM results in much higher yield and ultimate strength,while keeping the high elongation of the wrought material.Heat treatments above 950°C show no grain enlargement compared to the as built condition, but the cellular dendritic structure is completely dissolved. This results in lower hardness and yield strength compared to the as built condition.HT2 (heat treatment at 950°C) results in a promising heat treatment to be applied on 316L produced by SLM. HT2 reduces the yield and ultimate strength of the as built samples but564still resulting in higher values than requirement from EM10216-5:2013 standard (190MPa).

HT2 reduces the yield and ultimate strength of the as built samples but564still resulting in higher values than requirement from EM10216-5:2013 standard (190MPa). In addition, the absorbed energy, measured by the Charpy test,is significantly increased compared to the as build condition.Inclusions or formation of oxides can reduce and exhibit a wide spread in the elongation values of the SLM processed 316L.AcknowledgmentsThis research was supported by the ENGIE Research and Technology Division. The authors acknowledge ENGIE Research and Technology Division for the use of the SLM280HL machineReferences[1][2]P. Mercelis and J. Kruth, “Residual stresses in selective laser sintering and selective laser melting,” Rapid Prototyp. J., vol. 12, no. 5, pp. 254–265, Oct. 2006. I. Yadroitsev, A. Gusarov, I. Yadroitsava, and I. Smurov, “Single track formation in selective laser melting of metal powders,” J. Mater. Process. Technol., vol. 210, no. 12, pp. 1624–1631, Sep.

Gradl3 , Shamsaei1,2*Shuai Shao1,2, Nima1 National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn,AL 36849, USA2 Department of Mechanical Engineering, Auburn University, Auburn, AL36849, USA 3 NASA Marshall Space Flight Center, Propulsion Department, Huntsville, AL 35812, USACorresponding author: shamsaei@auburn.eduPhone: (334) 844-4839AbstractThis study examines and compares the effect of different heat treatments (HT) on the microstructure and mechanical properties of Monel K500 fabricated using laser powder bed fusion (L-PBF) and laser powder directed energy deposition (LP-DED) technologies. The as-fabricated Monel K500 specimens exhibited dendritic microstructure and elemental micro-segregation due to high cooling rates induced during the fabrication process.

The as-fabricated Monel K500 specimens exhibited dendritic microstructure and elemental micro-segregation due to high cooling rates induced during the fabrication process. The applicability of HT proposed in the literature for wrought Monel K500 was investigated for L-PBF and LP-DED using four different HT procedures involving hot isostatic pressing (HIP), solution annealing (SA), and aging. The mechanical properties of test specimens were evaluated using uniaxial tensile testing at room temperature. The microstructural evolution of test specimens during HT was analyzed using a scanning electron microscope. For all HT conditions investigated, L-PBF Monel K500 specimens consistently displayed higher strength and lower ductility compared to the LP-DED counterparts.

For all HT conditions investigated, L-PBF Monel K500 specimens consistently displayed higher strength and lower ductility compared to the LP-DED counterparts. The HT procedure involving HIP at 1160°C for 3hr at 100 MPa, SA at 1100°C for 15 min, and three step aging at 610°C for 16 hr, 540°C for 6 hr, and 480°C for 8 hr resulted in highest strength for both L-PBF and LP-DED fabricated Monel K500.Keywords: Additive manufacturing; Monel K500; Ni-base superalloys; microstructure; tensile propertiesIntroductionMonel K500 (UNS N05500) is a γ' precipitation hardened Ni-base superalloy and its major alloying element, copper (~30 wt.% Cu) provides solid solution strengthening and improved corrosion resistance [1– 3]. This alloy is particularly attractive for its high strength and hardness, and find applications in offshore and marine industries [4,5].

This alloy is particularly attractive for its high strength and hardness, and find applications in offshore and marine industries [4,5]. Due to the high hardness, conventional manufacturing and machining of Monel K500 alloy are often costly [6,7].Additive manufacturing (AM) can be an attractive alternative for fabricating net-shaped Monel K500 parts with little to no machining at reduced cost and time owing to its offered advantage of design flexibility and on-demand manufacturing [8,9]. Among different AM technologies, laser powder bed fusion (L-PBF) and laser powder directed energy deposition (LP-DED) are widely used [10]. Both processes fabricate parts in a layer-by-layer fashion with L-PBF process using scanning laser beams to selectively melt and fuse metal powder that is uniformly spread over a build plate and LP-DED process injecting metal powder into laser beam through a nozzle [11–13].

With AM being relatively new, the mechanical properties of additively manufactured Monel K500 parts have not been thoroughly investigated.Due to the unique thermal characteristics of AM processes, such as rapid cooling and solidification, the AM microstructure in the non-heat treated (NHT) condition are different from the wrought counterparts and often varies among different AM processes due to their difference in thermal history [14–16]. Moreover, the presence of AM process-induced volumetric defects can be detrimental to the mechanical properties of 361additively manufactured (AMed) parts [17]. This requires post-fabrication heat treatments (HT) to make the mechanical properties of AMed parts comparable to their wrought counterparts [18,19]. The existing HT procedures for the Monel K500 alloys are established for the wrought material and might not be optimum for L-PBF and LP-DED Monel K500.

The existing HT procedures for the Monel K500 alloys are established for the wrought material and might not be optimum for L-PBF and LP-DED Monel K500. This motivates investigating the effect of different HT on the mechanical properties of L-PBF and LP-DED Monel K500. Four different HT procedures following the literature for the wrought Monel K500 alloy are designed and their effects on the tensile properties of L-PBF and LP-DED Monel K500 are evaluated to establish a favorable HT procedure for the L-PBF and LP-DED Monel K500.Experimental proceduresThe chemical compositions of the two powder batches used for the fabrication of L-PBF and LP- DED Monel K500 specimens are presented in Table 1. For L-PBF, vertical near net-shaped specimens of 10 mm gage diameter and 80 mm length were fabricated by Quadrus Corporation using a Concept Laser M2 L- PBF machine and for LP-DED, vertical cylindrical bars of 15 mm diameter and 100 mm length were fabricated by RPM Innovations Inc.

All specimens underwent hot isostatic pressing (HIP) at 1163°C for 3 hours at a pressure of 100 MPa and furnace cooled (FC) to room temperature, followed by solution annealing (SA) at 1100°C for 15 min, and argon quenched (AQ) to room temperature. Finally, the hot isostatic pressed and annealed specimens were divided into four different groups and subjected to single and multi-step aging cycle, as listed in Table 3.Table 3. Thermal treatments applied for the Monel K500 specimens in this study, designated as HT1, HT2, HT3, and HT4.DesignationsHT1HT2HT3HT4HIP1163°C/100MPa/3hr FCSA1100°C/15min AQAging: Step 1 610°C/8hr AQ 610°C/16hr AQ 610°C/16hr FC 610°C/16hr FCAging: Step 2480°C/8hr AQ 540°C/6hr FCAging: Step 3480°C/8hr AQSmall coupons were excised from the NHT and heat treated Monel K500 specimens and microstructural characterizations were performed on the transverse plane, i.e., parallel to the build direction.

The coupons were mounted on cold epoxy resin and the top surface were mirror polished to remove any surface scratches according to the ASTM-E3 standard [20]. A Zeiss Crossbeam 550 scanning electron362microscope, equipped with electron backscatter diffraction (EBSD), backscattered electron (BSE), and energy dispersive X-ray spectroscopy (EDS) detectors, was used for microstructure characterization.The heat treated specimens were machined to final geometry (see Figure 1) and polished to remove any surface roughness prior to the uniaxial tensile testing at room temperature according to the ASTM E8 standard [21]. Uniaxial tensile testing was performed at a strain rate of 0.005 mm/mm/min. Due to the travel limit, the extensometer was attached to the specimens during the initial stage of tensile tests up to a strain value of 0.015 (mm/mm) to calculate the yield strength. Upon removal of the extensometer, the tests were continued in force-controlled mode till complete fracture.

Upon removal of the extensometer, the tests were continued in force-controlled mode till complete fracture. For all HT conditions, at least three specimens were tested and their average values were reported.Figure 1. (a) Small coupon for microstructural characterization (plane parallel to the build direction was characterized) and (b) the geometry of the tensile specimens used in this study (Dimensions are in mm).Results and discussionBSE micrographs obtained on the transverse plane of the L-PBF and LP-DED fabricated Monel K500 samples are presented in Figure 2 for the NHT and heat treated conditions. The NHT condition showed dendritic microstructure for both L-PBF and LP-DED Monel K500 (see Figure 2(a) & (b)) and elemental micro-segregation (see Figure 3 for LP-DED Monel K500), which are typical for the AMed metallic parts due to high cooling rates during fabrication.

Upon full HT, the dendritic microstructure was dissolved into the matrix as solid solution for all HT conditions (see Figure 2(c) - (j)).363Figure 2. BSE micrographs taken from the transverse plane of L-PBF and LP-DED samples: in (a) & (b) NHT, in (c) & (d) HT1, in (e) & (f) HT2, in (g) & (h) HT3, and in (i) & (j) HT4 conditions. The arrow indicates the build direction.364Figure 3. (a) Elemental maps and (b) X-ray spectrum obtained by EDS from the transverse plane of NHT LP-DED Monel K500.The inverse pole figure (IPF) maps obtained from the EBSD analysis on the transverse plane of the L-PBF and LP-DED samples are presented in Figure 4 for the NHT and heat treated conditions. For NHT and heat treated conditions, the LP-DED samples showed coarser grains compared to the L-PBF ones. In the NHT and heat treated LP-DED samples, the majority of the grains had sizes above 100 µm, while in L-PBF samples, finer microstructure with below 100 µm grain sizes were observed.

In the NHT and heat treated LP-DED samples, the majority of the grains had sizes above 100 µm, while in L-PBF samples, finer microstructure with below 100 µm grain sizes were observed. For all HT conditions, both L- PBF and LP-DED samples exhibited recrystallization and subsequent grain growth with presence of annealing twins in the microstructure (see Figure 4).365Figure 4. IPF maps taken from the transverse plane of L-PBF and LP-DED samples: in (a) & (b) NHT, in (c) & (d) HT1, in (e) & (f) HT2, in (g) & (h) HT3, and in (i) & (j) HT4 conditions. Material points are colored according to the normal direction of the TD plane. Arrow indicates the build direction.BSE micrographs obtained from the transverse plane of L-PBF and LP-DED samples are presented in Figure 5 for different HT conditions. The nano-sized γ' precipitates appeared to form in all the HT conditions (see Figure 5). Samples that undergo additional aging steps such as HT3 and HT4, formed coarser γ' precipitates.

The nano-sized γ' precipitates appeared to form in all the HT conditions (see Figure 5). Samples that undergo additional aging steps such as HT3 and HT4, formed coarser γ' precipitates. This is mainly due to longer soaking time during aging, which promotes growth of the precipitates [22,23]. In particular, the size of γ' precipitates in HT4 is larger than those in HT3 due to longer soaking time (6 hr longer) associated with an additional aging step.366Figure 5. BSE micrographs taken from the transverse plane of L-PBF and LP-DED specimens: in (a) & (b) NHT, in (c) & (d) HT1, in (e) & (f) HT2, in (g) & (h) HT3, and in (i) & (j) HT4 conditions. Red arrowheads point to γ' precipitates. Black arrow indicates the build direction.Engineering stress-strain curves and the corresponding ultimate tensile strength (UTS), yield strength (YS), and elongation at failure (%EL) for the different heat treated L-PBF and LP-DED fabricated Monel K500 tensile specimens are presented in Figure 6.

From HT1 to HT4, increasing trends in YS and UTS were observed for both L-PBF and LP-DED specimens, while decreasing trends were observed for ductility. The highest strength for HT4 could be attributed to larger γ' precipitates at higher volume fractions. For all HT conditions, L-PBF specimens showed higher strength and lower ductility compared to LP-DED counterparts. For instance, HT1 L-PBF Monel K500 specimens showed 37% and 19% higher YS and UTS respectively, and 12% lower %EL than HT1 LP-DED. Higher strength in L-PBF specimens can be attributed to the finer grain structures present in the microstructure compared to LP-DED process (see Figure 4). Among the heat treated L-PBF and LP-DED Monel K500 tensile specimens, HT4 L-PBF Monel K500 exhibited similar tensile properties compared to the wrought Monel K500 alloy [24] (see Figure 6 (b)).

Among the heat treated L-PBF and LP-DED Monel K500 tensile specimens, HT4 L-PBF Monel K500 exhibited similar tensile properties compared to the wrought Monel K500 alloy [24] (see Figure 6 (b)). To gain deeper insights into the impact of HT, ongoing research is focused on analyzing the fatigue performance of L-PBF and LP-DED Monel K500 alloy.367Figure 6. Tensile behavior of HT wrought, L-PBF, and LP-DED Monel K500 (a) Engineering stress-strain curves (b) bar chart showing variations in tensile properties such as UTS, YS, and %EL.Both L-PBF and LP-DED Monel K500 tensile specimens showed ductile fracture with dimples present on the tensile fracture surfaces (see Figure 7). However, larger sized dimples were observed for the LP-DED specimens compared to L-PBF counterparts. The higher elongation observed in the LP-DED specimens can be attributed to the coarser grain size (see Figure 4).

The higher elongation observed in the LP-DED specimens can be attributed to the coarser grain size (see Figure 4). In such cases, the increase in elongation comes at the expense of strength (see Figure 6).368Figure 7. Tensile fracture surfaces of heat treated L-PBF and LP-DED Monel K500 specimens: in (a) & (b) HT1, in (c) & (d) HT2, in (e) & (f) HT3, in (g) & (h) HT4 conditions.ConclusionsIn this study, the effect of different heat treatments (HT) procedures was investigated on the microstructure and tensile properties of laser powder bed fused (L-PBF) and laser powder directed energy deposited (LP-DED) Monel K500. The following conclusions were drawn:1. Dendritic microstructure was observed for both L-PBF and LP-DED Monel K500 samples in the non-heat treated conditions, which resolved after HT.2. Heat treated L-PBF samples exhibited finer grain structure compared to LP-DED counterparts. 3.

Heat treated L-PBF samples exhibited finer grain structure compared to LP-DED counterparts. 3. HT condition with longer soaking time during aging showed higher strength in both L-PBF and LP- DED Monel K500, which was ascribed to the growth of the γ' precipitates during aging.4. L-PBF Monel K500 specimens showed higher strength compared to LP-DED ones, which was ascribed to the finer grain structure.5. HT4 offered the highest strength for both L-PBF and LP-DED Monel K500 specimens.AcknowledgmentThis study is partially supported by the National Aeronautics and Space Administration (NASA) under Cooperative Agreement No. 80MSFC19C0010. This paper describes objective technical results and analysis. Any subjective views or opinions that might be expressed in the paper do not necessarily represent the views of the NASA or the United States Government.369References[1]Z. Chen, C. Wang, C. Tang, Y.Z. Lek, S.Y. Kandukuri, H. Du, H. Gao, K.

viewed PaperTHE EFFECT OF HEAT TREATMENT ON THE MICROSTRUCTURE AND TENSILE PROPERTIES OF LASER POWDER BED FUSED (L-PBF) HAYNES 214Shaharyar Baig1,2, Shuai Shao1,2, Paul R. Gradl3, Nima Shamsaei1,2*1 Department of Mechanical Engineering, Auburn University, Auburn, AL 36849, USA 2 National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849, USA 3 NASA Marshall Space Flight Center, Propulsion Department, Huntsville, AL 35812, USACorresponding author: Email: shamsaei@auburn.edu Phone: (334) 844 4839AbstractThe as-fabricated microstructure of Haynes 214 by laser powder bed fusion (L-PBF) is dendritic due to the high cooling rates, which is generally considered undesirable and removed via heat treatments. In this study, the effects of various heat treatments on the microstructure and tensile properties of laser powder bed fused Haynes 214 are investigated. Test specimens were fabricated on a L-PBF system in the vertical orientation.

Test specimens were fabricated on a L-PBF system in the vertical orientation. Multi-step heat treatments were performed including stress-relief, hot isostatic pressing, solution annealing and ageing. Examining the microstructures revealed columnar grains in the non-heat treated (NHT) condition with grain growth occurring after stress relief. Room temperature tensile tests showed the lowest strengths in the NHT condition. The highest strengths were seen in the aged condition which was attributed to gamma prime precipitation hardening.

Room temperature tensile tests showed the lowest strengths in the NHT condition. The highest strengths were seen in the aged condition which was attributed to gamma prime precipitation hardening. Finally, the tensile fracture surfaces indicated ductile failure mode in both the NHT and heat treated conditions with the measured elongation being the highest for NHT and lowest in the aged condition.Keywords: Additive manufacturing (AM), laser powder bed fusion (L-PBF), Haynes 214, microstructure, tensile propertiesIntroductionNi-based superalloys are commonly used in oxidizing environments due to their ability to retain strength, even at elevated temperatures [1–3]. For γ- γ’ Ni-based alloys, the main strengthening results from the γ-solid solution consisting of various alloying elements as well as the formation of γ’ phase and grain boundary (GB) carbide precipitates at elevated temperatures [3–5].

In the case of Haynes 214 Ni-based superalloy, the addition of Cr and Al allows protective layers of oxides scales to form on the surface which provides resistance to carburizing, nitriding, and corrosion in harsh environments [6]. However, the high content of Al in this alloy typically makes it susceptible to strain-age cracking, especially during any post-process treatments, where the formation of coarse γ’ can lead to a stressed microstructure and makes conventional processing challenging [7–9].528Additive manufacturing (AM) looks to be a promising technology to fabricate functional parts as it minimizes the risk of introducing any defects from post-process treatments [9]. Moreover, AM allows the fabrication of complex assemblies, whereby several assembly parts can be fabricated into one integral component [10].

Moreover, AM allows the fabrication of complex assemblies, whereby several assembly parts can be fabricated into one integral component [10]. Due to their unique thermal history, the non-heat treated microstructures of AM parts are not only different from their wrought counterparts, but can also vary among different AM technologies [11]. Note that existing heat treatment (HT) schedules available in literature were established for wrought Haynes 214 [9,12]. These HT routines may not result in consistent microstructures and optimized mechanical properties in L-PBF Haynes 214 parts. As such, further studies are required to examine the effect of different heat treatments on the resulting mechanical properties and establish process-structure-property relationship for L-PBF Haynes 214.In this study, the microstructure and tensile properties of additively manufactured Haynes 214 processed through laser powder bed fusion (L-PBF) are examined.

The microstructures in the NHT as well as various solution annealed (SA) only and solution annealed + aged (SA + A) conditions were characterized to identify which conditions may be the most favorable in providing acceptable tensile properties. Room temperature tensile testing was then performed on a few selected conditions as well as the fracture mechanisms were investigated.Experimental ProcedureSpecimens were fabricated using argon-atomized Haynes 214 powder supplied by Praxair technologies with the chemical composition shown in Table 1. The build consisted of vertically oriented specimens processed using the core L-PBF parameters presented in Table 2. A near net- shape geometry was adopted for the as-fabricated samples as can be seen in Figure 1 (a), which would later be cut and/or machined into the final coupon and specimen geometries for microstructural and mechanical characterization, respectively.Table 1.

Chemical composition of Haynes 214 feedstock powder used for fabrication. YElementNiCrAlFeSiCoWt. %Bal.15.74.322.360.020.020.01Table 2. L-PBF Haynes 214 core process parameters used in this study.Laser Power (W) 200Laser speed (mm/s) 1000Layer Thickness (µm) 40Hatch distance (µm) 70Cylindrical microstructural coupons, 5 mm in thickness, were excised from the as- fabricated specimens. Some of the microstructural coupons were heat treated, as explained later in this section, to examine the microstructure under various conditions. The coupons were mounted in epoxy resin in the transverse direction (TD), to investigate the microstructure in the plane parallel to the build direction (BD), and ground and polished to a mirror finish. The remaining tensile specimens were surface machined, to reduce any effects of surface roughness resulting from the L-PBF process. The final tensile test specimen geometry is shown in Figure 1 (b).529Figure 1.

The final tensile test specimen geometry is shown in Figure 1 (b).529Figure 1. (a) As-fabricated specimen geometry and (b) final tensile test specimen geometry after machining.All stress relief (SR), solution anneal, and ageing treatments were performed in a Thermo Scientific box furnace. Various multi-step heat treatments were investigated and a list of all heat treatment steps with their associated temperatures, times, and cooling methods are summarized in Table 3. All heat treated specimens were stress relieved in the first step at 1095°C for 2h followed by a hot isostatic pressing (HIP) at 1163°C/100 MPa for 3.5h. The HIP specimens were then solution annealed at 1100°C for 2h as the next step and some of the specimens underwent a final heat treatment step of ageing at 788°C for 8h.

The HIP specimens were then solution annealed at 1100°C for 2h as the next step and some of the specimens underwent a final heat treatment step of ageing at 788°C for 8h. Microstructural characterization was performed between each heat treatment step, and three conditions were selected for tensile specimens, namely: NHT, SR+HIP+SA1100 (which is simplified as the SA condition from hereon) and SR+HIP+SA1100+A788 (simplified as the SA + A condition from hereon).Table 3. Heat treatment conditions with their respective temperatures, pressures, times, and cooling methods. Temperature/Pressure/Time 1095°C/Ambient/2h 1163°C/100MPa/3.5h 1100°C/Ambient /2h 788°C/Ambient /8hHeat treatment step Stress Relief (SR) Hot Isostatic Press (HIP) Solution Anneal (SA) Age (A)Cooling method Furnace cool Furnace cool Water quench Water quenchOptical microscopy was performed on polished coupon cross sections using a Keyence VHX 6000 for porosity measurements.

The Keyence software was used for analysis of the optical micrographs by setting the thresholds to allow the overall porosity content as well as the size and the geometric features for the defects to be measured. For microstructural characterization, the grain size and morphology were evaluated using a Zeiss Crossbeam 550 scanning electron microscope (SEM) equipped with an electron backscatter diffraction (EBSD) detector. A backscattered electron (BSE) detector was also used to collect images for further microstructural characterization using the SEM.530Tensile testing was performed on an MTS servohydraulic load frame with a load cell capacity of 100 kN. Room temperature tensile testing was conducted according to ASTM E8 under displacement control mode at a strain rate of 0.01 s-1. An extensometer was attached to the specimens for the first 0.035 mm/mm of strain in order to measure the modulus and 0.2% yield strength.

An extensometer was attached to the specimens for the first 0.035 mm/mm of strain in order to measure the modulus and 0.2% yield strength. For each condition, at least 2 tests were performed and the average values for all of the properties measured were reported. Tensile tests were concluded upon fracture and all fractured specimens were carefully preserved to measure the elongations and reduction in areas as well as to examine the fracture surfaces.Results and DiscussionThe average density, based on optical defect analysis, was measured as 99.88% for the NHT condition. After SR and the HIP cycle, no significant changes were seen in the density value, which was measured to be 99.881%. The histogram of defect size distribution is presented in Figure 2 which shows there were no defects measured larger than 30µm in the HIP condition.

The histogram of defect size distribution is presented in Figure 2 which shows there were no defects measured larger than 30µm in the HIP condition. This indicates that HIP was able to shrink some of the larger defects however did not significantly affect some of the smaller gas-entrapped pores which were still largely present and observed for both NHT and the specimens which underwent the HIP cycle.Figure 2. Porosity size distribution for the NHT and HIP specimens.Figure 3 (a)-(d) presents the inverse pole figure (IPF) plots in the transverse plane for the various heat treated conditions. The average grain sizes for the NHT, SR, SA and SA+A were measured as 24.4 µm, 22.7 µm, 25.4 µm and 24.8 µm, respectively. It can be observed in Figure 3 (a) that in the NHT condition, the grains are columnar and oriented in the build direction. This type of grain structure can be attributed to the epitaxial grain growth enabled by the repeated, the L-PBF process.

This type of grain structure can be attributed to the epitaxial grain growth enabled by the repeated, the L-PBF process. Such layer-by-layer melting-solidifying process, characteristic microstructures may not be desirable in load bearing applications as it leads to mechanical anisotropy. After SR, the grain morphology changes significantly with the grains becoming moreto531equiaxed as shown in Figure 3 (b). The presence of equiaxed grains as well as the higher volume fraction of high angle grain boundaries (i.e., average misorientation angles >10° measured as 40% for NHT and 80% for SR) indicates that significant grain growth has occurred when the specimens were subjected to elevated temperatures of 1065°C during the SR treatment. No further changes were observed in the grain size and morphology after the SA and SA+A treatments as shown in Figure 3(c)-(d).Figure 3. IPF plots in the TD for (a) NHT; (b) SR; (c) SA and (d) SA+A conditions.Figure 4.

IPF plots in the TD for (a) NHT; (b) SR; (c) SA and (d) SA+A conditions.Figure 4. presents the BSE micrographs in the transverse plane for the NHT, SA and SA+A specimens. A dendritic microstructure was observed in the NHT condition which results from fast solidification under the high cooling rates inherent to the L-PBF process. After subsequent SR followed by SA at 1100°C for 2h and water quenching, this structure dissolved, and only Al-rich phases (suspected to be topologically close-packed (TCP) intermetallics) could be seen as shown in Figure 4 (d). Heating the specimens to above the solvus temperature of γ’ (i.e., 955°C) allowed any precipitates which may have formed during the SR step, to dissolve into the Ni-based matrix resulting in a γ solid solution. The final ageing step at 788°C for 8h resulted in the formation of fine γ’ precipitates as well as GB carbides as presented in Figure 4 (e)-(f).

The final ageing step at 788°C for 8h resulted in the formation of fine γ’ precipitates as well as GB carbides as presented in Figure 4 (e)-(f). It is important to note that the rate of cooling during each heat treatment is critical in determining the amount of age- hardening observed, as γ’ can precipitate in Haynes 214 in the temperature range of 595-925°C [9]. Interestingly, no cracks were observed in the SEM micrographs for either the NHT, SA or SA+A conditions. This can be attributed to the Al content for the feedstock powder measured at 4.32 wt.%, which is below the 4.5 wt.% commonly believed to make these alloys non-weldable and thus affecting their processability through AM [13].532Figure 4. BSE micrographs taken in the TD for (a)-(b) NHT; (c)-(d) SA and (e)-(f) SA+A conditions.A summary of the measured tensile properties is presented in Figure 5.

BSE micrographs taken in the TD for (a)-(b) NHT; (c)-(d) SA and (e)-(f) SA+A conditions.A summary of the measured tensile properties is presented in Figure 5. The elastic moduli were measured as 165, 220 and 221 GPa for the NHT, SA and SA+A specimens, respectively, indicating that significant changes in the phase compositions occurred during heat treatments. In the NHT condition, the yield and ultimate strengths were measured as 516 MPa and 714 MPa, respectively, with a significant amount of necking observed at failure. This is also reflected in Figure 6 (a) which shows a large reduction in the cross-sectional area of the gauge section for the NHT specimen. SR followed by SA resulted in solid solution strengthening which improved the ultimate tensile strength (Su) to 934 MPa, however no significant increase in the yield strength (Sy) was observed, which was measured as 524 MPa.

After the ageing treatment, both the yield and the ultimate tensile strengths increased to 618 MPa and 1033 MPa, respectively, which was also accompanied by a reduction in the measured elongation. This can be attributed to the formation of fine γ’ precipitates during the ageing process which improved the strengths, however the presence of GB carbides may have led to a reduction in the ductility, as they can serve as early crack initiation sites.533Figure 5. Summary of the tensile yield strength (Sy), ultimate tensile strength (Su), and the elongation to failure (% Elongation) for the NHT, SA and SA+A conditions.The SEM fracture surface images for the NHT, SA and SA+A conditions are presented in Figure 6. In the NHT condition, large pores and dimples can be observed on the fractured surface. The reduction in area was measured to be 84% which correlates with the significant amount of plastic deformation during necking of the specimens.

The reduction in area was measured to be 84% which correlates with the significant amount of plastic deformation during necking of the specimens. In comparison, for the SA and SA+A conditions, some cracks as well as dimples were observed on the fractured surfaces. For these two conditions, the reduction in areas were not as high as the NHT condition, however were still significant at 58% and 51% for the SA and SA+A conditions, respectively. Dimple fracture was seen in all of the conditions with a typical cup and cone fracture shape. This indicates that the fracture modes were ductile for all conditions with the tensile strengths being governed by the amount of solid solution strengthening as well as precipitation hardening from the respective heat treatments.534Figure 6.

SEM images of fracture surfaces for the (a)-(b) NHT; (c)-(d) SA1100 and (e)- (f) SA+A conditions.ConclusionsIn this study, the effect of various heat treatments on the microstructure and mechanical properties of L-PBF Haynes 214 were investigated. A comprehensive microstructural characterization was carried out for various multi-step heat treatments which included SR, HIP, SA and SA+A. Grain growth was observed after SR treatment and no significant changes seen in the grain size/morphology upon subsequent heat treatments. BSE micrographs revealed a dendritic structure in the NHT condition, which dissolves after solution treatment. Fine γ’ precipitates and grain boundary carbides started to form in the SA+A condition which resulted from the ageing step at 788°C for 8h. The highest tensile strengths were reported in the SA+A condition which was attributed to the formation of strengthening precipitates.

The highest tensile strengths were reported in the SA+A condition which was attributed to the formation of strengthening precipitates. The tensile fracture surfaces revealed that ductile failure occurred in all conditions with the highest elongation being measured for the NHT and the lowest in the SA+A condition.535AcknowledgmentThis research is partially supported by the National Aeronautics and Space Administration (NASA) under Cooperative Agreement No. 80MSFC19C0010. This paper describes objective technical results and analysis. Any subjective views or opinions that might be expressed in the paper do not necessarily represent the views of the NASA or the United States Government.References[1]Pollock TM, Tin S. Nickel-based superalloys for advanced turbine engines: Chemistry, microstructure, and properties. Journal of Propulsion and Power 2006;22:361–74. https://doi.org/10.2514/1.18239.[2]Sajjadi SA, Zebarjad SM.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of Heat Treatment on the Microstructures of CMSX-4® Processed through Scanning Laser Epitaxy (SLE)Amrita Basak1, and Suman Das1, 21George W. Woodruff School of Mechanical Engineering, 813 Ferst Dr NW, Georgia Institute of Technology, Atlanta, GA 30332 2School of Materials Science and Engineering, Georgia Institute of Technology, 771 Ferst Dr NW, Atlanta, GA 30313Engineered components made of superalloys routinely undergo heat treatment procedures to tailor their microstructures and properties. In this study, the effect of heat treatment on the microstructures and microhardness of single-crystal nickel-base superalloy CMSX-4® is investigated. Samples of CMSX-4® are fabricated using scanning laser epitaxy (SLE), a laser (LPBF)-based additive manufacturing process.

Samples of CMSX-4® are fabricated using scanning laser epitaxy (SLE), a laser (LPBF)-based additive manufacturing process. Microstructural powder bed characterizations of the as-deposited and the heat treated CMSX-4® samples are performed using optical microscopy, scanning electron microscopy, x-ray diffraction, and Vickers microhardness measurements. The results show that the microstructure is homogenized with reductions in the eutectic volume fraction after heat treatment. The microhardness values are also improved upon heat treatment. This work is sponsored by the Office of Naval Research through grant N00014-14- 1-0658.fusionKeywords: Additive Manufacturing, Scanning Laser Epitaxy (SLE), Nickel-base, Superalloys, CMSX-4®, Heat-treatmentIntroductionThe performance of gas turbines improves with increase in the operating temperature creating a hostile environment for the hot-section components.

Currently, nickel-base superalloys are used to fabricate hot-section components as they offer higher yield strength with increase in operating temperatures primarily due to the presence of the secondary  precipitate phases that form an antiphase boundary and lock the associated dislocation movements. However, due to the rotational movements of hot-section components under very high temperature, they suffer from material losses due to thermal abrasion, corrosion, and oxidation. With several hundred such components installed in each gas turbine engine, their total replacement cost at the time of overhaul is upwards of hundreds of thousands or even millions of dollars per engine.

Hence, there is a great commercial interest in developing additive manufacturing (AM)-based repair processes that are capable of restoring the parent metallurgical microstructure and geometry at damage locations.Scanning laser epitaxy (SLE) is a laser powder bed fusion (LPBF)-based AM process that creates equiaxed (EQ), directionally solidified (DS) and single-crystal (SX) structures of nickel superalloys onto like-chemistry substrates. In SLE, a galvanometer controlled laser beam is focused upon the powder bed that generates a melt-pool. The melt-pool travels along the length of the substrate fusing the completely melted powder to the partially melted substrate. Under the proper operating conditions, the solidified microstructure has been demonstrated to follow the morphology of the underlying substrate [1].

Under the proper operating conditions, the solidified microstructure has been demonstrated to follow the morphology of the underlying substrate [1]. The SLE process has shown significant potential for193one-step repair of CMSX-4® [2-4], René N5 [5, 6], René 142 [7], MAR-M247 [8], René 80 [9], and IN100 [10, 11].CMSX-4® is a second generation ultrahigh-strength SX superalloy developed using a multidimensional approach over a ten year period through the joint efforts of Canon-Muskegon and Allison at the beginning of the 1990s [12]. The alloy demonstrates a significant improvement of 35 oC in the turbine airfoil temperature capability compared to the first-generation SX superalloys such as CMSX-2® and CMSX-3® [13]. CMSX-4® is derived from CMSX-2® and enjoys the strengthening effects of rhenium (Re) (3 wt. %) like other second generation SX superalloys. Re partially segregates to the  matrix, and thereby, delays coarsening of the phase [14].

%) like other second generation SX superalloys. Re partially segregates to the  matrix, and thereby, delays coarsening of the phase [14]. CMSX-4® is strengthened by the solid-solution-strengthening effects of chromium (Cr), tungsten (W), and tantalum (Ta) and the precipitation-hardening effects of aluminum (Al) and titanium (Ti) that form the phase [12].Although CMSX-4® has excellent thermo-mechanical properties, heat treatment via aging (below 1090 oC) or solutioning (above 1090 oC) and aging is usually applied to further improve the mechanical properties. Heat treatment is instrumental in reducing elemental segregation. Heat treatment procedures also refine the dispersion of phases and decrease the amount of  eutectics. Additionally, carbide precipitates are also redistributed during heat treatment and re- precipitate at the grain boundaries [8]. In the present work, SLE deposited CMSX-4® samples were subjected to heat treatment via aging.

In the present work, SLE deposited CMSX-4® samples were subjected to heat treatment via aging. In order to relieve the residual stresses and enable precipitation of the strengthening phases, the samples were heated and kept at 1080 oC for an hour. Thereafter, the samples were cooled at 5.6 oC /minute or faster to a temperature below 900 oC. Samples were then given a typical commercial low-temperature precipitation heat-treat cycle consisting of 4 hours at 900 oC, followed by air cooling at minimum average cooling rate of 5.6 oC / minute to 650 oC to fully develop the strengthening phases. The microstructures of the as- deposited and the heat-treated samples were analyzed using optical microscopy (OM) and scanning electron microscopy (SEM), and the results are reported.Materials and MethodsPowder of nickel-base superalloy CMSX-4® produced by the Praxair Surface Technologies through argon gas atomization was used in this study. The composition of the powder is reported in Table I.

The composition of the powder is reported in Table I. The powder was deposited on investment-cast CMSX-4® substrates. The morphology and the cross-section of powder were analyzed using OM and SEM. The CMSX-4® powders were mostly spherical, although some irregularity in shape was found as shown in Fig. 1(a). The powder was mounted in Bakelite and polished to a mirror finish. The cross-section was analyzed under an optical microscope for the inspection of internal porosity, and no major internal porosity was detected as shown in Fig. 1(b).Table I. Chemical composition of the CMSX-4® powder (wt. %)CMSX-4®Cr Co Mo Re W Al Ti Ta Ni 5.6 1.0 6.5 Bal 6.59.70.4 3.0 6.4194Porosity(a)(b)200 m100 mFig. 1. (a) SEM image of the CMSX-4® powder and (b) OM image of the CMSX-4® powder cross- section after polishing.SLE processed samples were sectioned along the length and width for the inspection of the microstructure. A Buehler automated saw was used to section the samples.

A Buehler automated saw was used to section the samples. The first half was retained while the second half was subjected to heat treatment. Each section was mounted in Bakelite and polished to a mirror finish; starting with 80 grit paper and progressively increasing the size to 1200 grit. The samples were then rough-polished using 5 µm and 3 µm diamond solutions. Finally, the samples were fine polished using a 0.5 µm colloidal alumina suspension. The polished samples were then etched with Marble’s reagent (50 ml HCl, 50 ml H2O, and 10.0 gm CuSO4) to eliminate the phase and reveal the dendritic microstructure. A Leica DM6000 optical microscope was thereafter used to take the images. The microstructural investigation of the SLE processed CMSX-4® was carried out on a Hitachi SU8230 SEM. The SEM was equipped with an Oxford Instruments’ Aztec Energy EDX system which was used for elemental analysis.

The SEM was equipped with an Oxford Instruments’ Aztec Energy EDX system which was used for elemental analysis. Vickers microhardness measurements were carried out using a Buehler microhardness indenter with a fixed load at 2000 gf. The hardness values were extracted in Vickers microhardness scale (HV). Details on hardness measurements may be found elsewhere [11].SLE Experimental ProcedureThe SLE process was conducted on rectangular investment-cast SX CMSX-4® substrates having dimensions of 38.10 mm x 11.85 mm x 2.54 mm. Each substrate was placed into a 38.10 mm x 11.85 mm recess cut into an IN625 base plate. The CMSX-4® powder was placed above the substrates using rectangular wells cut into an Aluminum mask plate. Once the samples were prepared, they were placed into an atmospheric process chamber that was purged with high purity (99.999%) Argon.

Once the samples were prepared, they were placed into an atmospheric process chamber that was purged with high purity (99.999%) Argon. A 1kW Ytterbium fiber laser (IPG Photonics, Model: YLS-1000) was used with a Cambridge Technologies galvanometer scanner to focus the beam on top of the substrate to a Gaussian beam diameter of 40 m. A raster scan pattern across the width of the sample generated a melt pool that linearly propagated along the length of the substrate. All the samples were run with 25.4 m scan spacing.195Results and DiscussionFig. 2 shows the length-wise cross-sectional microstructure of the CMSX-4® as-deposited sample. The average deposit thickness was found to be more than 1500 m in a single-pass. As shown in Fig. 2, very few pores were observed. The deposit was also crack-free and dense. The deposit region showed a finer microstructure compared to the substrate region as shown in Fig. 2.

As shown in Fig. 2, very few pores were observed. The deposit was also crack-free and dense. The deposit region showed a finer microstructure compared to the substrate region as shown in Fig. 2. The microstructural refinement might be attributed to the localized heating and fast cooling due to the high-speed laser heat source.SubstrateDepositFig. 2. Representative transverse OM image of the as-deposited CMSX-4® sample showing crack- free and dense deposit. The black line represents the substrate-deposit interface.200 mDeposit(a)(b)SubstrateFig. 3. OM image showing interface microstructure (a) of the as-deposited sample, and (b) the heat-treated sample. The black line represents the substrate-deposit interface.The zoomed view of the substrate-deposit interface region is shown in Figs. 3(a) and 3(b) for the as-deposited and the heat-treated samples. No significant difference in the interface was observed near the interface region as shown in Figs. 3(a) and 3(b).

3(a) and 3(b) for the as-deposited and the heat-treated samples. No significant difference in the interface was observed near the interface region as shown in Figs. 3(a) and 3(b). The microstructure of as- deposited CMSX-4® revealed three types of dendritic microstructures such as the [001] columnar dendrites, [100] columnar dendrites, and the equiaxed dendrites. The [001] columnar dendritic microstructure grew epitaxially from the partially re-melted grains of the substrate as shown in196Fig. 3(a). The microstructure of the deposited CMSX-4® retained the columnar structure after the heat treatment as shown in Fig. 3(b).The major constituents of the SLE processed microstructure of CMSX-4® were the  matrix, the  precipitates in the  matrix, and the eutectics. Figure 4(a) illustrates an SEM image of the deposit region showing the major constituents in an as-deposited sample. Figure 4(b) illustrates the major constituents in the SLE deposited CMSX-4® after heat treatment.

Figure 4(b) illustrates the major constituents in the SLE deposited CMSX-4® after heat treatment. The eutectic volume fraction is visibly reduced after the heat treatment process as shown in Figure 4(b). However, after the heat treatment additional phases were found to be formed in the deposit region as shown in Fig. 4(b).(a)(b)Eutectic ’10 mTCP PhasesEutectic ’Fig. 4. SEM image showing morphologies of various constituents of SLE processed CMSX-4® deposit – (a) as-deposited sample and (b) heat-treated sample.500 nm5 m’(a)(b)(c)’’500 nmFig. 5. Representative SEM image showing /morphology in the (a) substrate region, (b) deposit region of the as-deposited sample, and (c) deposit region of the heat-treated sample.As shown in Fig. 5(a), the primary precipitates showed a characteristic dimension in the range of 2 m in the substrate region whereas in the deposit region, the precipitates in the range of 50 nm as shown in Fig. 5(b).

5(b). The deposit region shows approximately 25% reduction in the197size for the as-deposited sample. Microstructural refinement in the deposit region is visible from Figs. 5(a) and 5(b). Fig. 5(c) illustrates the size distribution in the heat-treated sample. The primary size is of order 100 nm in the heat-treated samples. The heat-treated CMSX-4® showed predominantly cuboidal precipitates as compared to polygonal shaped precipitates in the as- deposited sample as shown in Figs. 5(b) and 5(c).To compare the size of the primary precipitates, their equivalent diameters were calculated. The equivalent diameter is defined as D= (A/)0.5, where A is the area of a precipitate determined by measuring the perpendicular sides of an individual precipitate. The average equivalent diameter, <D>, is calculated by the mean of ~100precipitates. Fig.

The average equivalent diameter, <D>, is calculated by the mean of ~100precipitates. Fig. 6 illustrates the particle size distribution (PSD) with the theoretical Lifshitz-Slyozov-Wagner (LSW) [8] distribution superimposed for the as-deposited and heat-treated CMSX-4® samples. The normalized diameter  is defined as D/<D>. The SLE processed PSD of the γ' precipitate differs considerably from the theoretical LSW distribution. This might be due to the fact that the size of the  precipitates in the interdendritic regions was found to be larger than those present within the dendrite region. Note that while performing this calculation, the γ/precipitates in the eutectic pools were not considered.

Note that while performing this calculation, the γ/precipitates in the eutectic pools were not considered. The size showed a closer match to the theoretical PSD for the heat-treated sample.Probability Distribution Function (PDF)Calculated PDF of the Heat-Treated CMSX-4Calculated PDF of the As-Deposited CMSX-4LSW Probability Distribution Function (PDF)0.000.050.100.150.200.25Fig. 6. Particle size distribution (PSD) of the  precipitates – as-deposited and heat-treated sample.Fig. 7(a) shows SEM images of secondary phase precipitates that were formed after the heat treatment. These precipitates were identified as topologically-closed-pack (TCP) phases. The size of these precipitates were of order 500 nm. These precipitates were observed to line up along the grain boundary surrounded by the phase. The elemental EDS maps of these particles are shown in Figs. 7(b) through 7(d).

These precipitates were observed to line up along the grain boundary surrounded by the phase. The elemental EDS maps of these particles are shown in Figs. 7(b) through 7(d). The precipitates were found to be rich in W, Re, and Cr.198(a)(b)(c)(d)WReCrAdditional Phase5 mFig. 7. (a) Representative SEM image showing morphology of the secondary phase precipitate in the CMSX-4® deposit after the heat treatment. (b)- (d) EDS Map of the precipitate.WEDCEDCDCDCEDCEDCEDCNiReCoEETiDCEEDCDCDCDCEDCDCCrEWE10 mMoDCReAlEECoEDCEDCNiECrTaTiTaDCSEM ImageESEM ImageAlMo(a)(b)EDCEFig. 8.

8(a) and 8(b) illustrate EDS maps of a representative region in the CMSX-4® deposit before and after the heat treatment, respectively. Note that the dendritic core and the eutectic regions are marked by DC and E, respectively. Ti and Ta showed segregation to the eutectics. W, Al, and Re showed segregation to the core (Figs. 8(a) and 8(b)). Other elements did not show any strong segregation behavior as shown in Figs. 8(a) and 8(b).In order to quantify the elemental segregation, EDS line scans were performed across representative eutectic regions in the deposit region before and after the heat treatment,200respectively (Figs. 9(a) and 9(b)). The line profiles of a representative element (Cr) across a eutectic region in the deposit before and after the heat treatment are shown in Figs. 9(c) and 9(d), respectively.

9(a) and 9(b)). The line profiles of a representative element (Cr) across a eutectic region in the deposit before and after the heat treatment are shown in Figs. 9(c) and 9(d), respectively. From the elemental line profiles, the maximum and the minimum counts were extracted and a segregation parameter was defined as degree of segregation (k) = maximum cps/minimum cps (where cps denotes counts per second). The k values for various elements are plotted in Fig. 9(e) for the deposit region before and after the heat treatment. Although, in comparison to the as-deposited sample, the eutectic phase showed increased elemental segregation for the heat-treated CMSX-4®, the dimension of the eutectic pools was significantly smaller in the heat-treated sample.The hardness of an alloy is a strong function of the size, distribution, and volume fraction and the residual stresses generated during the laser processing.

In the present study, microhardness results indicated a hardness value in the range of 430-500 HV in the deposit region. The heat- treated CMSX-4® showed 10% increase in the hardness values as reported in Table II. This increase may be attributed to the uniform distribution obtained after the heat treatment. The yield stress was approximated as, y ~ HV/3 (HV measured in GPa) [8]. Table II shows the calculated yield stress for the as-deposited and the heat-treated CMSX-4® samples [8].Table II. Summary of microhardness results for CMSX-4®.Sample TypeAs-Deposited CMSX-4®Region Substrate Interface DepositHV2000 400 423 464HV (GPa) 3.92 4.15 4.55y (GPa) 1.30 1.38 1.51Heat-Treated CMSX-4®Substrate Interface Deposit405 446 4923.97 4.37 4.831.32 1.46 1.61ConclusionsIn the present study, SLE process was demonstrated to deposit dense crack-free deposits of CMSX-4® exceeding 1500 m in a single-pass. Commercial heat treatment suited for CMSX-4® was performed.

Commercial heat treatment suited for CMSX-4® was performed. A detailed study was conducted to characterize the microstructural changes due to the heat treatment. The microstructures of CMSX-4® became uniform after the heat treatment with more than 80% reductions in the eutectic volume fraction. The heat treatment also improved the microhardness values in the deposit region. Future investigations will be carried out to characterize TCP phase formations in CMSX-4®. The effect of heat treatment on the texture and grain size distribution will also be analyzed.Acknowledgments and DisclosuresThis work is sponsored by the Office of Naval Research through grant N00014-14-1-0658. Dr. Suman Das is a cofounder of DDM Systems, a start-up company commercializing SLE technology. Dr. Das and Georgia Tech are entitled to royalties derived from DDM Systems’ sale of products201related to the research described in this paper. This study could affect their personal financial status.

This study could affect their personal financial status. The terms of this arrangement have been reviewed and approved by Georgia Tech in accordance with its conflict of interest policies.References[1] Basak A, Das S. Epitaxy and Microstructure Evolution in Metal Additive Manufacturing. Annual Review of Materials Research 2016;46:125-49. [2] Acharya R, Bansal R, Gambone JJ, Das S. A Coupled Thermal, Fluid Flow, and Solidification Model for the Processing of Single-Crystal Alloy CMSX-4 Through Scanning Laser Epitaxy for Turbine Engine Hot-Section Component Repair (Part I). Metallurgical and Materials Transactions B 2014;45:2247-61. [3] Acharya R, Bansal R, Gambone JJ, Das S. A Microstructure Evolution Model for the Processing of Single-Crystal Alloy CMSX-4 Through Scanning Laser Epitaxy for Turbine Engine Hot-Section Component Repair (Part II). Metallurgical and Materials Transactions B 2014;45:2279-90. [4] Basak A, Acharya R, Das S.

Various materials have been successfully manufactured via the MBJ process; however, appropriate post- process heat treatments are required to enhance their mechanical performance as compared to the wrought or other additively manufactured counterparts. This study aims to investigate the effect of post-manufacture heat treatment on the microstructure and mechanical properties of MBJ 17-4 PH stainless steel (SS). Various heat treatment procedures following the standard routes for the wrought 17-4 PH SS are conducted to evaluate their effects on the tensile behavior of MBJ 17-4 PH SS. The mechanical behavior of the MBJ 17-4 PH SS in various heat treatment conditions is discussed based on their corresponding microstructure.Keywords: Metal binder jetting (MBJ); Stainless steel; Microstructure; Heat treatment; TensileIntroductionMetal binder jetting (MBJ) is categorized as a multi-step additive manufacturing (AM) process following ISO/ASTM52900-15 [1].

However, contrary to the laser-based powder bed AM processes, MBJ is a non-fusion AM technique. In this technique, the spread powder is selectively joint with typically a polymeric binder, layer-by-layer until the desired part geometry is formed in a box of powder. The parts are further consolidated by several stages as follows and shown in Figure 1:Curing or drying: the box produced (green parts and powder together) is placed in the oven to dry the ink. This will allow the green parts to harden -thanks to binder crosslinking reaction- to be recovered and handled. • Powder removal: during which the non-amalgamated powder is extracted from the box in order to recover the parts. Moving the parts at this stage may be difficult because they are in green condition.

• Powder removal: during which the non-amalgamated powder is extracted from the box in order to recover the parts. Moving the parts at this stage may be difficult because they are in green condition. The powder is further recovered and recycled.879Debinding: this operation consists of removing organic binder from the part by using evaporation and decomposition mechanisms and leaving only a metal part with high porosity of about 35-40 %. • Sintering: it gives the part its final shape. The part is heated to 80-90 % of the melting temperature of the material to sinter and coalescence the powder particles.

• Sintering: it gives the part its final shape. The part is heated to 80-90 % of the melting temperature of the material to sinter and coalescence the powder particles. Porosity is reduced, and the density is increased due to a fairly uniform shrinkage fairly about 14-16 % (following the definition of shrinkage with sintering indicated in ISO 4492), so that the part proportions are maintained.The MBJ technique is compatible with any powder material in contrast to other fusion- based AM techniques (e.g., laser beam powder bed fusion (LB-PBF), laser beam directed energy deposition (LB-DED), laser-foil-printing (LFP) [2], etc.). Since the fabrication of the part via MBJ is at a low temperature (typically 80°C in our case), there are no such issues as oxidation, and residual stress, and the powder surrounding the part in the box is highly recyclable. No support structure is needed for the MBJ parts in contrast to fusion-based techniques.

No support structure is needed for the MBJ parts in contrast to fusion-based techniques. However, depending on the shape of the part, a sintering setter may be needed; this setter that is printed at the same time as the part is not physically connected to it as the support structure for the LB-PBF. The setter is used to prevent any part walls collapsing where there are overhangs or to support compacting of the material during sintering in order to prevent any unwanted deformations.Figure 1. Typical Metal Binder Jetting (MBJ) process cycle [3].There are drawbacks to the MBJ technique; it is a multi-step process to finish a part ready to be deployed in any application, as-sintered parts usually have densities around (97-98%), which is lower than that of LB-PBF specimens but with the same range as Metal Injection Molding ones. Due to the substantial shrinkage during sintering, the main challenge in MBJ would be controlling the size and the mass of parts.

Due to the substantial shrinkage during sintering, the main challenge in MBJ would be controlling the size and the mass of parts. Difficulty in controlling the tolerances for the size and mass of the parts increases with component size. The larger the part is, the more difficult it will be to maintain tight tolerances and avoid cracking during sintering.880Various material systems such as nickel-based superalloys [4–6], titanium alloys [7,8], and iron and stainless steels [9] have been manufactured via the MBJ process, and efforts have been made to optimize the sintering temperature and post heat treatment processes for different materials. Mostafaei et al. [5] optimized the sintering and post process heat treatment for the MBJ Inconel 625 to be comparable to the conventionally manufactured counterparts. Similarly, Karlsson et al.

Mostafaei et al. [5] optimized the sintering and post process heat treatment for the MBJ Inconel 625 to be comparable to the conventionally manufactured counterparts. Similarly, Karlsson et al. [10]studied the post heat treatment effects on the mechanical properties of MBJ AlCoCrFeNi high entropy alloy and reported superior mechanical strength and ductility as compared to those of Inconel 625 and Inconel 718 alloys.One of the prominent materials used in the AM industry is 17-4 PH stainless steel. This material possesses a desirable combination of strength, ductility, fatigue strength, and high corrosion resistance, which finds use in key engineering, defense, and energy sectors. Therefore, there has been a strong desire to adopt 17-4 PH SS for AM to take advantage of its benefits [11]. The 17-4 PH SS has been successfully manufactured via various laser-based AM techniques.

Therefore, there has been a strong desire to adopt 17-4 PH SS for AM to take advantage of its benefits [11]. The 17-4 PH SS has been successfully manufactured via various laser-based AM techniques. There have been several studies that investigated the effect of powder recycling [12,13], shielding gas type [11,14], build orientation [15,16], surface roughness, and heat treatment [17] on the microstructure and mechanical properties (tensile and fatigue) of LB-PBF 17-4 PH SS.The main challenge with the AM 17-4 PH SS is to find the most appropriate heat treatment process to enhance its mechanical properties considering the detrimental impacts of induced defects by AM processes. It has been shown that although the peak-age heat treatment (CA-H900) increases the tensile strength of LB-PBF 17-4 PH SS, the fatigue performance is deteriorated after applying this heat treatment [18].

This is attributed to the presence of defects induced by AM process, which makes the high strength material more sensitive in crack initiation and fatigue failure. Nezhadfar et al. [17] investigated the effect of various heat treatment procedures on the microstructure and mechanical properties (tensile and fatigue) of LB-PBF 17-4 PH SS. They reported the CA-H1025 heat treatment procedure as an appropriate heat treatment results in a good combination of tensile strength, ductility, and fatigue performance.To the best of the authors’ knowledge, there is not much literature found on MBJ 17-4 PH SS. This study aims to investigate the effect of various heat treatments on the microstructure and mechanical properties of MBJ 17-4 PH SS. Different heat treatment procedures are applied on the MBJ 17-4 PH SS specimens. The effect of solution heat treatment and aging temperature and duration is studied on the microstructure of MBJ 17-4 PH SS specimens.

The effect of solution heat treatment and aging temperature and duration is studied on the microstructure of MBJ 17-4 PH SS specimens. Eventually, the tensile properties are obtained for each and every one of the applied heat treatments, and the results are discussed according to their corresponding microstructure.Experimental ProceduresThe specimens in this study were fabricated using 17-4 PH SS powder with a particle-size distribution D10=7 µm, D50 = 14 µm, D90 = 23 µm. The chemical composition is listed in Table 1.Table 1. Chemical composition of 17-4 PH SS powder used for fabricating the MBJ specimens.(Wt.

The chemical composition is listed in Table 1.Table 1. Chemical composition of 17-4 PH SS powder used for fabricating the MBJ specimens.(Wt. %)C 0.05Cr 16.00Ni 4.50Cu 4.40Mn 0.30Si 0.70Nb 0.24881Fe Bal.The specimens were printed on a Digital Metal Machine DMP2000 with the followingprocess parameters.Layer thickness: 42µm, the powder is spread in the Y direction - Build strategy: “Lissel-4PH” from Digital Metal, with a printing direction in X-axis - Build orientation of the specimens in the box: X-axis - Cure time and duration: the box is cured a 205°C for 36 hoursThe build layout is shown in Figure 2 including the cubic coupons (10x10x10 mm3) for microstructure characterization and cylindrical bars (diameter of 12 mm, length of 80 mm), which were further machined to tensile test specimens. The green parts were gone through an air debinding cycle of soaking at 345 °C for 2 hours after powder removal.

The green parts were gone through an air debinding cycle of soaking at 345 °C for 2 hours after powder removal. The sintering cycle was further conducted on the specimens using an Elnik MIM3015 T furnace under pure hydrogen at 1380 °C for 2 hours. After the sintering process, the specimens were undergone hot isostatic pressing (HIP) at 1160 °C for 3 hours at a pressure of 1020 bars. The porosity of the HIPed specimens was measured via image analysis to be 0.01%.Figure 2. Build layout for fabricating MBJ 17-4 PH SS specimens.Cubic coupons shown in Figure 3 were fabricated for microstructure investigation (see Figure 2 as well). The cube coupons were cut transversely to characterize the microstructure in XY-plane (i.e., perpendicular to the build direction and parallel to the loading direction seen in Figure 3(b). Specimens were polished using sandpapers with 320-4000 grits, and then mirror- finished using a Chemo-met pad along with using 0.05 µm colloidal silica suspension.

Specimens were polished using sandpapers with 320-4000 grits, and then mirror- finished using a Chemo-met pad along with using 0.05 µm colloidal silica suspension. The specimens were further polished using a vibratory polisher with 0.02 µm colloidal silica suspension for 2 hours prior to conducting any microstructure analysis. A Zeiss 550 Crossbeam scanning electron microscope (SEM) with an Oxford electron backscatter diffraction (EBSD) detector was used for microstructure characterization.882Figure 3. (a) The cube coupon for microstructure characterization (XY-plane parallel to load direction was characterized), and (b) the tensile geometry following ASTM E8.Horizontal cylindrical bars with 10 mm diameter and 67mm length were fabricated via the MBJ process to be machined into the tensile specimens with the geometry following the ASTM E8 as shown in Figure 3(b). The specimens were heat treated in an argon-purged box furnace following the procedures listed in Table 2.

The specimens were heat treated in an argon-purged box furnace following the procedures listed in Table 2. The heat treatment procedures are following the ASTM A693 standard [19]; however, in some cases, the solution heat treatment and aging steps were modified to evaluate new heat treatment procedures for this material. The tensile tests were further carried out at a 0.001 s-1 strain rate to evaluate the mechanical properties of the MBJ 17-4 PH SS specimens in various heat treatment conditions.Table 2.

The tensile tests were further carried out at a 0.001 s-1 strain rate to evaluate the mechanical properties of the MBJ 17-4 PH SS specimens in various heat treatment conditions.Table 2. Heat treatment procedures applied on the HIPed MBJ 17-4 PH SS specimens.Heat Treatment (HT) HT1 (CA-H900) HT2 (Mod* CA-H900) HT3 (Mod1 CA-H1025) HT4 (Mod2 CA-H1025) *Mod = ModifiedProcedure 1050 °C/0.5 hour/Air cooled → 482 °C/1 hour/Air cooled 1100 °C/1 hour/Air cooled → 482 °C/1 hour/Air cooled 1100 °C/1 hour/Air cooled → 552 °C/0.5 hour/Air cooled 1100 °C/1 hour/Water cooled → 552 °C/0.5 hour/Air cooledResults and DiscussionThe microstructure characterization results for the as-sintered and HIPed conditions obtained by EBSD are presented in Figure 4(a) and (b), respectively. As seen in Figure 4(a), sintering was achieved successfully and parts were fabricated with an adequate density. However, there are still huge pores remained in the microstructure.

As seen in Figure 4(a), sintering was achieved successfully and parts were fabricated with an adequate density. However, there are still huge pores remained in the microstructure. In addition, the grain boundaries miss- orientation reveals a high fraction (~67%) of low angle grain boundaries (i.e., 5°-15°). This shows that the grain structure has not developed completely and further post-process thermal treatments are required. The microstructure in the as-sintered condition consists of a martensitic matrix with the presence of coarse strings of δ-ferrite with sharp edges.It can be observed in Figure 4(b) that conducting the HIP cycle closes the pores and results in an approximately full dense structure. The microstructure still consists of martensite along with δ-ferrite. However, HIP makes most of the δ-ferrite grains finer and thinner strings.

The microstructure still consists of martensite along with δ-ferrite. However, HIP makes most of the δ-ferrite grains finer and thinner strings. Moreover, a883higher fraction of high-angle grain boundaries is formed after performing HIP as compared to the as-sintered condition, which witnesses the occurrence of recrystallization of the grain during HIP. Although the HIP process is performed in the austenitic region, there is δ-ferrite phase retained in the microstructure. This may be due to the chemical composition of the powder used for fabrication as well as considering the low cooling/heating rates and low solidification rate during the sintering of the parts. Some δ-ferrite may be formed during heating above the 1280 °C up to 1380 °C (i.e., the sintering temperature), and also some liquid phase will form during sintering which the low solidification rate results in solidification in the order of liquid→ δ-ferrite+austenite→ δ- ferrite+martensite [20].

The δ-ferrite retained in the microstructure from the sintering process will remain in the structure even by performing high temperature heat treatment; however, the morphology of the grains may be changed from sharp edge string-like grains to more circular grains.Figure 4. EBSD image quality (IQ) juxtaposed by grain boundaries (GB) showing the microstructure of MBJ 17-4 PH SS in (a) as-sintered, and (b) HIPed conditions.The microstructure of MBJ 17-4 PH SS after different heat treatments (see Table 2) is presented in Figure 5. An overall observation of the microstructures indicates the martensitic microstructure along with the distributed δ-ferrite phase in the microstructure. As was mentioned, even performing solution heat treatment at the austenite region could not resolve the δ-ferrite in the microstructure; however, the morphology and the grain sizes varied according to the applied heat treatment.

Figure 5(a) shows the microstructure of MBJ 17-4 PH SS specimen after HT1 (see Table 2) including the martensite and δ-ferrite grains. As seen, conducting solution heat treatment at 1050 °C for 0.5 hours could slightly re-arrange the δ-ferrite grains as compared to the HIP condition (see Figure 4); the δ-ferrite grains after heat treatment broken apart and are not as continuous as the ones in the HIP microstructure. However, there are still stretched string-like δ- ferrite grains in the microstructure.Increasing the solution heat treatment temperature from 1050 °C/0.5 hr to 1100 °C/1 hr with the same following aging procedure (i.e., 482 °C/1 hr) in HT2 could break the δ-ferrite grains to smaller and more circular ones seen in Figure 5(b). Higher solution heat treatment temperature for a longer duration can result in more fraction of recrystallized austenite grains and result in the re-arrangement of δ-ferrite grains in the microstructure.

As a result of a higher degree of recrystallization in the microstructure can be referred to the higher high angle grain boundaries in HT2 condition as compared to the HT1.884Figure 5. EBSD IQ+GB and inverse pole figure (IPF) maps of MBJ 17-4 PH SS for: (a) HIP+HT1, (b) HIP+HT2, (c) HIP+HT3, and (d) HIP+HT4 specimens.It can be observed for the HT3 condition shown in Figure 5(c) that at the constant solution heat treatment temperature and duration (i.e., 1100 °C/ 1 hour) as that of HT2, increasing the aging temperature to 552 °C can result in a slight growth of δ-ferrite phase, and higher low angle grain boundaries as compared to the HT2 condition. Higher fraction of low angle boundaries is due to885the high temperature aging; however, not sufficient time for the rearrangement of dislocations to form high angle grain boundaries [21].

Higher fraction of low angle boundaries is due to885the high temperature aging; however, not sufficient time for the rearrangement of dislocations to form high angle grain boundaries [21]. In the HT4 procedure, specimens were quenched in water after solution heat treatment with the same temperatures to investigate the effect of quenching media on the microstructure and mechanical properties as compared to the HT3 condition. It can be observed in Figure 5(d) that the grain structure is not changed considerably as compared to the HT3 counterpart; however, due to rapid cooling, the δ-ferrite grains are thinner than the HT2 condition, which is ascribed to finer lath martensite in the microstructure of HT4 as compared to HT3.Tensile properties of the MBJ 17-4 PH SS specimens are presented in Figure 6. As seen, the as-sintered specimen has very low ductility due to the presence of large pores (see Figure 4(a)).

As seen, the as-sintered specimen has very low ductility due to the presence of large pores (see Figure 4(a)). However, the ductility of the material increases significantly after performing HIP ascribed to the closure of the pores (see Figure 4(b)), which enhances the structural integrity of the material. Although HIP increases the ductility of the material, it does not change the strength of the material and the tensile strength (i.e., Sy and Su) are comparable with that of the as-sintered specimen. A similar observation has been reported for the LB-PBF 316L SS in which the reduction in the porosity by preheating the build platform increases the ductility without any impact on the strength of the material [22]. Conducting the heat treatment after HIP, however, influences both the strength and ductility of the MBJ 17-4 PH SS specimen.

Conducting the heat treatment after HIP, however, influences both the strength and ductility of the MBJ 17-4 PH SS specimen. HT1 increases the tensile strength and reduces the ductility as compared to the HIP condition, yet possesses higher ductility than as-sintered specimen. The high strength in the HT1 specimen is attributed to the formation of nano-size Cu- enriched precipitates in the microstructure [23]. Considering the presence of the δ-ferrite phase, known as a brittle phase with the limited active slip systems during deformation [24], the MBJ 17- 4 PH SS specimen shows a brittle behavior result in relatively low ductility.The presence of δ-ferrite retained from the solidification is inevitable in all the heat treatment conditions, however, conducting prior heat treatment may alter its morphology and size.

As seen in Figure 5(b), by conducting HT2 with higher solution heat treatment temperature and duration (i.e., 1100 °C/1hr) as compared to that of HT1 (i.e., 1050 °C/0.5hr), the δ-ferrite grains are broken apart and become more circular with less sharp edges compared to those in HT1 condition. This change in the morphology of the brittle δ-ferrite phase by applying HT2 makes the material less brittle and enhances the ductility of the material compared to the HT1 specimen as confirmed in Figure 6.Furthermore, at a constant solution heat treatment temperature as HT2, increasing the aging temperature to higher temperature and shorter duration (i.e., 552 °C/0.5 hr) in HT3, the ductility of the material increases considerably, and the strength decreases. It can be seen in Figure 5(c) that although most δ-ferrite grains have more curved edges, the grain growth occurred due to the higher aging duration, which can be the reason for the reduction in strength of the material.

At the same time, it is expected that the ductility decreases due to the presence of larger δ-ferrite grains in the microstructure. However, since HT3 is an averaging heat treatment condition, it is reported that the Cu-enriched precipitates are coarsened and semi-coherent with the matrix. In this case, dislocations cannot shear the precipitates and must loop around and bypass the precipitates following the Orowan mechanism [25], which results in higher ductility as compared to the HT2 condition. It has been well established that conducting overage heat treatment on the AM 17-4 PH SS decreases the strength and increases the ductility to some extent [17].886Figure 6. Tensile properties (i.e., Sy, Su, and El%) of MBJ 17-4 PH SS specimens in various conditions.

Tensile properties (i.e., Sy, Su, and El%) of MBJ 17-4 PH SS specimens in various conditions. The standard deviation is shown by error bars.According to the effect of morphology and size of δ-ferrite phase, as well as the size of the Cu-enriched precipitates on strength and ductility, a heat treatment is required to result in preferably small and more circular δ-ferrite phase, or at least with curved edges, as well as not coarsened Cu-enriched precipitates. This combination may enhance the strength as compared to the HT3 condition without sacrificing much ductility. As seen, changing the quenching media from air to water in the HT4 condition increases the strength and slightly decreases the ductility as compared to the HT3 condition.

As seen, changing the quenching media from air to water in the HT4 condition increases the strength and slightly decreases the ductility as compared to the HT3 condition. Interestingly enough, the strength obtained by HT4 is comparable with that of HT2, slightly lower than that of HT1, yet the ductility of the HT4 specimen is higher than that of both HT2 and HT1 specimens.ConclusionsIn this study, the effect of heat treatment on the microstructure and tensile properties of MBJ 17-4 PH SS was investigated. The microstructure of MBJ 17-4 PH SS specimens was characterized in the as-sintered condition and after HIP as the reference. Further, various heat treatment conditions with different solution heat treatment and aging temperatures and duration, as well as different quenching media were applied to obtain acceptable strength and ductility for the MBJ 17-4 PH SS. The following conclusions are drawn based on the experimental results:1.

The following conclusions are drawn based on the experimental results:1. There were large pores and a high fraction of low angle grain boundaries remained in the microstructure of the as-sintered specimen, which the HIP process successfully closed all the pores and developed the grain structure having a higher fraction of high angle grain boundaries. This was attributed to the occurrence of grain recrystallization.8872. The strings of δ-ferrite phase with sharp edges retained in the microstructure of as fabricated MBJ 17-4 PH SS specimens, which was attributed to the chemical composition of material as well as low cooling/heating rates and the low solidification rate of MBJ process after sintering in a very high temperature in which the δ-ferrite was formed. 3. Conducting well-known peak age heat treatment (i.e., CA-H900) called HT1 herein, increased the strength considerably, while it decreased the ductility.

3. Conducting well-known peak age heat treatment (i.e., CA-H900) called HT1 herein, increased the strength considerably, while it decreased the ductility. This was attributed to the formation of nano-size Cu-enriched precipitates.4. Increasing the solution heat treatment temperature and duration (e.g., HT2 vs. HT1) broke the δ-ferrite phases apart to more circular in shape with more curvature in edges. This was found to enhance the ductility of the material.5. Comparable tensile strength (i.e., Sy, and Su) and higher ductility than the peak-age condition (i.e., HT1) were obtained by changing the solution heat treatment quenching media to water.References[1]ISO/ASTM 52900:2015(E), Standard Terminology for Additive Manufacturing – General Principles – Terminology, (n.d.).[2] C.-H. Hung, W.-T. Chen, M.H. Sehhat, M.C.

Various heat treatment procedures (CA-H900 and CA-H1150) were conducted to evaluate their effects on the tensile behavior of LP-DED 15-5 PH SS. Scanning electron microscopy was used to characterize the microstructural features and the fracture surfaces. Tensile tests were performed to evaluate the mechanical properties at cryogenic and room temperatures. The mechanical behavior of the LP-DED 15-5 PH SS in different heat treatment conditions is discussed based on their microstructures and fracture surfaces. Reduction in area of CA-H1150 treated specimens after tensile tests was significantly higher than CA-H900 ones, while the ultimate tensile and yield strengths of CA-H900 specimens were higher compared to the CA-H1150 ones.

The CA-H1150 specimens had relatively coarser Cu-enriched precipitates, and exhibited greater necking compared to CA-H900 specimens.Keywords: Laser powder directed energy deposition (LP-DED), precipitation hardened stainless steel, heat treatment effects, cryogenic temperature, tensile fractureIntroduction15-5 PH stainless steel (SS) is a martensitic stainless steel known for its excellent mechanical properties and corrosion resistance at various temperatures [1,2]. This alloy is typically used to manufacture components in aerospace industries, nuclear reactors, and medical applications where the service temperature ranges from cryogenic to elevated [3]. Since components used in such sectors often possess intricate geometries which are difficult and traditional subtractive manufacturing, additive manufacturing (AM) offers a viable solution.

Since components used in such sectors often possess intricate geometries which are difficult and traditional subtractive manufacturing, additive manufacturing (AM) offers a viable solution. One of the AM techniques is laser powder-directed energy deposition (LP-DED), in which a laser beam is used to melt the powder delivered by an inert carrier gas through nozzles.time-consumingto be manufactured viaHowever, the mechanical behaviors of parts manufactured by AM techniques differ from those fabricated via traditional manufacturing techniques due to residual stresses and process-induced defects such as lack of fusion, keyholes, and gas-entrapped pores [4–7]. In addition, AM parts tend to have different microstructures than wrought ones, which often remain even after applying the same thermal post-processing techniques [8–10].

In addition, AM parts tend to have different microstructures than wrought ones, which often remain even after applying the same thermal post-processing techniques [8–10]. While a few research articles investigated the effects of build orientations, heat treatments and surface treatments on microstructural features and mechanical properties of additively manufactured 15-5 PH SS [2,11–18], there still exists a knowledge gap regarding the effect of heat treatments on tensile properties and the tensile fracture mechanisms at different temperatures. This study aims to investigate the effects of heat treatments on the microstructure and tensile behaviors of LP-DED 15-5 PH SS at cryogenic and room temperatures.372Experimental DetailsArgon atomized 15-5 PH SS powder supplied by Praxair Surface Technologies was used to fabricate the cylindrical bars (diameter = 11mm) by LP-DED technique in RPM Innovations (RPMI) 557 machine.

The detailed chemical composition of this powder is listed in Table 1. The cylindrical bars were fabricated using a process parameter set, laser power of 1070 W, layer height of 0.38 mm, laser scan speed of 1016 mm/min, and powder feed rate of 15.2 g/min under the argon environment.After fabrication, the cylindrical bars were stress-relieved at 650°C for an hour which was followed by hot isostatic pressing (HIP) at a temperature and pressure of 1163°C and 103MPa for 3 hours, respectively. The bars underwent solution treatment at 1050°C for half an hour (i.e., Condition A (CA)). After that, the specimens were divided into two groups: one was aged at 482°C for an hour (i.e., H900), known as the peak-age condition, and the other was over-aged at 621°C for 4 hours (i.e., H1150).

Both non-heat treated (NHT) and heat treated LP- DED 15-5 samples were cut at the gage section in the planes parallel to (i.e., longitudinal plane) and perpendicular to (i.e., radial plane) the build direction (BD). These samples were mounted and polished for microstructural characterization. A Zeiss Crossbeam 550 scanning electron microscope (SEM) was used to characterize the microstructure of the samples in both radial and longitudinal directions. The fully heat treated specimens were then machined to final geometries, as shown in Figure 1 for tensile tests.The machined LP-DED 15-5 PH SS specimens were subjected to quasi-static monotonic tensile tests on MTS servo-hydraulic testing machines with 100kN load cells. The tensile specimens were loaded at a strain rate of 0.005mm/mm/min at temperatures of -196°C and 21°C. These tests were conducted following the ASTM E8- 21 standard for room temperature and the ASTM E21-20 standard for cryogenic temperature.

These tests were conducted following the ASTM E8- 21 standard for room temperature and the ASTM E21-20 standard for cryogenic temperature. Extensometers were installed in the gauge section of test specimens to measure the strain. For cryogenic tests, the specimens and grips of the machine were submerged in liquid nitrogen and soaked for 5 minutes after the specimens were cooled down to -196°C. The specimens were then loaded to failure. After completion of the tests, the fractured specimens were removed from a liquid nitrogen container. The fracture surfaces were cleaned in an ultrasonic cleaner and were analyzed using an SEM.Table 1. Chemical composition of 15-5 PH SS powder used in this study.Elements Weight %Fe Cr Ni Cu Mn Si Nb O Mo N C P S OtherBal. 14.99 5.24 3.5 0.49 0.04 0.2 0.03 0.01 0.01 0.01 0.006 0.003 0.1373Figure 1.

14.99 5.24 3.5 0.49 0.04 0.2 0.03 0.01 0.01 0.01 0.006 0.003 0.1373Figure 1. Tensile specimen geometry of LP-DED 15-5 PH SS (All dimensions are in mm).Results and DiscussionThe inverse pole figure (IPF) maps obtained after electron backscatter diffraction (EBSD) analysis on the longitudinal planes of NHT, CA-H900, and CA-H1150 samples are shown in Figure 2. These EBSD analyses also showed the presence of a minimal fraction of retained austenite in the martensitic matrix. The fraction of retained austenite in NHT, CA-H900, and CA-H1150 samples was 0.1%, 0.1%, and 0.14%, respectively. Due to the high cooling rate during the AM process, the austenitic phase gets transformed into martensite. The formation of retained austenite might be because of the diffusion of austenite stabilizing elements (i.e., Cu, Ni) into the grain boundaries, which may have acted as nucleation sites for austenite grains [19].

The slightly higher fraction of retained austenite in the CA-H1150 sample could be because of the longer aging duration at a temperature close to the austenite reversion temperature [19].Figure 2. IPF maps obtained from the longitudinal planes of a) NHT b) CA-H900 and c) CA-H1150 samples of LP-DED 15-5 PH SS.The elemental maps obtained after energy dispersive X-ray spectroscopy (EDS) analysis on CA-H900 and CA-H1150 samples are presented in Figure 3 and Figure 4, respectively. Nano-sized Cu-enriched precipitates can be observed in the elemental maps of the CA-H900 sample, whereas the Cu-enriched precipitates were coarser in CA-H1150 samples. In addition to Cu-enriched precipitates, Cr-, Mn-, Mo-, and Nb-enriched precipitates were observed in both CA-H900 and CA-H1150 samples. Moreover, significant segregation of Ni can be observed in the CA-H1150 sample.374Figure 3. Elemental maps obtained from EDS analysis on a CA-H900 sample.Figure 4.

However, upon over-aging, these precipitates coarsen. The coarsened Cu-enriched precipitates lose coherency with the matrix, decreasing the alloy’s strength375[1,20]. The tensile strengths of both CA-H900 and CA-H1150 specimens were higher at the cryogenic temperature than at room temperature. At such a low temperature, the kinetic energy is not sufficient enough to exceed the activation energy required for the dislocation motion.[10,21].This reduced dislocation mobility makes plastic deformation difficult [3,22].Figure 5. Bar plots showing UTS, YS and RA at cryogenic and room temperature for CA-H900 and CA-H1150 specimens.The fracture surfaces for CA-H900 and CA-H1150 specimens tested at cryogenic and room temperatures are shown in Figure 6. The fracture surfaces of the CA-H900 specimens tested at cryogenic temperature had multiple facets and showed no necking, indicating transgranular brittle fracture.

The fracture surfaces of the CA-H900 specimens tested at cryogenic temperature had multiple facets and showed no necking, indicating transgranular brittle fracture. The CA-H900 specimens tested at room temperature exhibited slight necking, dimples, and facets on the fracture surfaces indicating a mixed- mode fracture. The fracture surfaces of the CA-H1150 specimen tested at cryogenic temperature also had features similar to CA-H900 specimens tested at the room temperature. However, at room temperature, the CA-H1150 specimens exhibited significant necking compared to the ones tested at -196°C. The presence of larger dimples on the fracture surface suggests more ductile fracture at room temperature for CA-H1150 specimens. In general, the CA-H1150 specimens exhibited greater extent of necking with larger dimples on the fracture surfaces when compared with the CA-H900 specimens.376Figure 6.

In general, the CA-H1150 specimens exhibited greater extent of necking with larger dimples on the fracture surfaces when compared with the CA-H900 specimens.376Figure 6. Tensile fracture surfaces of CA-H900 and CA-H1150 specimens at (a,b,e,f) -196°C and (c,d,g,h) 21°C.377ConclusionsThis study investigated the effect of heat treatments on the tensile properties of LP-DED 15-5 PH SS atcryogenic and room temperature. The following conclusions can be drawn from this study:Both non-heat treated and heat treated LP-DED 15-5 PH SS possessed martensitic microstructure with a minimal fraction of retained austenite. The fraction of retained austenite was higher in CA-H1150 samples due to longer aging at a temperature closer to austenite reversion temperature.After full heat treatments, Cu-enriched nano precipitates were observed in the martensitic matrix.

The specimens treated with CA-H1150 had coarser Cu-enriched precipitates compared to CA-H900 specimens due to longer aging at a temperature closer to austenite reversion temperature.The CA-H1150 specimens exhibited lower tensile strength and higher ductility than the CA-H900 specimens, which could be because of coarsened Cu-enriched precipitates.CA-H900 specimens exhibited brittle fracture behavior at cryogenic temperature, whereas the fracture mechanism switched to ductile-brittle mode at room temperature. In contrast, CA-H1150 specimens showed ductile-brittle mode fracture behavior at cryogenic temperature and ductile fracture behavior at room temperature.AcknowledgementThis study is partially supported by the National Aeronautics and Space Administration (NASA) under Cooperative Agreement No. 80MSFC19C0010. This paper describes objective technical results and analysis.

fect of In-Plane Voiding on the Fracture Behavior of Laser Sintered PolyamideDavid K. Leigh, David L. Bourell, Joseph J. Beaman Jr. The University of Texas at Austin, USA Harvest Technologies, Inc., USAARTICLE INFO Keywords: Additive Manufacturing (AM) Rapid Polyamide Selective Laser Sintering (SLS)ABSTRACT The primary contributors to poor mechanical properties in polyamide materials used during Selective Laser Sintering® are qualified. Methods to quantify the decreased mechanical properties, including Scanning Electron Microscopy (SEM) of fracture surfaces, are compared against each other and against mechanical properties of components fabricated using multiple process parameters.

Of primary interest are Ultimate Tensile Strength (UTS) and Elongation at Break (EOB) of tensile specimens fabricated under conditions that produce varying degrees of ductile and brittle fracture.INTRODUCTIONAdaptation of Laser Sintering for End-Use PartsLaser SinteringAdditive manufacturing is a family of relatively new manufacturing processes utilizing integration techniques of digital solid models and the layer-by-layer addition of materials to create a three-dimensional solid. Selective Laser Sintering®, developed at The University of Texas at Austin, utilizes three major steps to create the three- dimensional solid. The first step of the process is the computational evaluation the solid model and the subsequent creation of a set of cross-sectional vectors referred to as “slices,” with each slice having a field of x-y vectors that represent the cross-sectional area of the model.

Each slice is prepared based upon a predetermined “layer thickness” which is typically 0.004” to 0.010”. The second step of the process is to preheat a bed of powder; commonly a polyamide (nylon) material with an average particle size of 50 microns. Once the powder is heated to a sufficient temperature and the slice file is created, a directed energy beam (CO2 laser) scans the vector field on the prepared bed of powder. The fused layer reproduces the cross- section generated by the vector field (or slice). Once fused, a layer of powder equal to the predetermined layer thickness is added, heated, and subsequently fused to prior layers to create a solid physical model, illustrated in Figure 1.The laser sintering process has become accepted as the most robust additive manufacturing process and has been deemed suitable for the production of end-use parts in a number of applications.

This novel technique has allowed for streamlined design, decreased time to productize, minimal cost on engineering changes, and the ability to design for function rather than manufacturability. While there may be many benefits for certain applications, there are signficant challenges that must be addressed prior to a broad acceptance of this technology. The key challenge is the limited amount of theoretical or experimental science specific to use of laser sintering as a manufacturing process.Process Variability and Limitation of Mechanical PropertiesA significant limitation to any layer-based additive manufacturing process is the anisotropic nature of the process. In this case, a single layer (typically 0.004” thick) is fused for each layer. The properties within this layer are fairly consistent, but the bond between layers is typically viewed as weak.

In this case, a single layer (typically 0.004” thick) is fused for each layer. The properties within this layer are fairly consistent, but the bond between layers is typically viewed as weak. This is quantified through the placement of standard ASTM tensile specimens oriented in the X-axis (left to right), Y-axis (front to back), and Z-axis (vertical build direction). It is typically observed that there is little to no variation between the X and Y axes, but that the Z-axis will yield a significantly lower elongation at break (EOB) than is observed in the X and Y axes.The primary cause of the lower mechanical properties in the Z-axis can be attributed to layer-to-layer adhesion. There are several causes of this effect, but little has been done to illustrate what is actually happening. To illustrate the effect of material processing on mechanical properties, several techniques are used.

To illustrate the effect of material processing on mechanical properties, several techniques are used. As to the cause of poor adhesion between layers, there are several factors that may contribute to this:Large layer size – The powder insulates the thermal energy and prevents the energy of the beam from penetrating to the prior layer.Fig 1: Selective Laser Sintering® ProcessPoor powder quality – Used powder will yield a higher melt-flow and higher molecular weight. This is primarily caused through thermal aging and cross-linking of the polymer. The result of the higher melt-flow is a polymer that is slow to melt and prone to the creation of voids.Poor thermal control – If the powder bed is processed in a way that the temperature is too low or is cooled at a high rate, it will inhibit the formation of a melt pool that penetrates the previous layer.

The result can be weak layer-to-layer adhesion as well as dimensional distortions.Low laser energy – High scan speed, low laser power, and large spacings between scan vectors can lower the overall laser energy delivered to the powder bed. This decreased energy will not yield sufficient layer-to-layer adhesion.Formation of Layers in Laser SinteringFigure 2 is a graphical representation of the raw material powder bed. While the average particle size of the polyamide material is 50 microns, there is a range of fine particles in a randomly packed bed. This powder bed is preheated to a point just below the melting temperature. If the particles are too hot, they will pool and there will be no distinction between the scanned area and the powder bed. If the particles are too cold, they will not have sufficient internal energy to allow for a full melt.

If the particles are too cold, they will not have sufficient internal energy to allow for a full melt. In addition, a region heated primarily though laser energy will have signficant dimensional distortion as the fused layer rapidly cools and shrinks. Subsequent layers that are fused will try to shrink, creating residual stresses due to adhesion with previous layers.Figure 2: Cross Section of a Laser Sintering Powder BedOnce the powder bed has been pre-heated, the laser will scan the subject cross-section using a field of x-y vectors at a prescribed speed, spacing, and laser power. These settings are optimized to yield an ideal melt pool such that there is sufficient energy to have a full melt without losing precision. If the laser energy applied is too great, detail and precision are lost (analogous to writing on tissue paper with a large permanent marker.) Figure 3 illustrates the melting of the powder bed surface by the laser.

There is a corresponding depth of penetration directly proportional to the laser energy applied.Figure 3: Melt Pool Generated on a Powder BedEXPERIMENTAL PROCEDUREDesign of Experiments to Evaluate and Correlate Physical and Visual PropertiesThe objective of this research is to correlate the mechanical properties observed with visual inspection to obtain a better understanding of the failure mechanisms of laser sintered parts. The material was 3D Systems Duraform® PA, a Nylon 12 polyamide. Tensile specimens were added to a production build to evaluate the surface of a tensile specimen and the fracture surfaces of both X and Z direction tensile specimens. The process settings and build location of test specimens were such that the three failure modes typical for laser sintering would be present for this experiment:1. Delamination – A failure that is charactarized by very weak layer-to-layer adhesion.

Delamination – A failure that is charactarized by very weak layer-to-layer adhesion. The fracture of a z-axis oriented tensile specimen will break exclusively in the region between layers. Delaminated Z-direction tensile specimens were thermal parameters with laser power at less than 50% of recommended parameters. Brittle Fracture – Failure typically charactarized with ultimate tensile strength and offset yeild strength comparable to stated material datasheets, yet with a significant reduction in elongation at break. Processing paramters are what is typical for most prototype builds.3. Ductile Fracture – This is the desired failure mode. A ductile fracture will yield nearly isotropic parameters. Processing paramaters have been optimized for build quality with a sacrifice in efficiency and cost.Observation and Quantification of Actual Tensile Specimen FailureTensile properties were measured using an MTS Insight 10 at Harvest Technologies.

The specimens produced were ASTM D638 specimens with a cross-sectional area at fracture of roughly 0.125” X 0.500”. The procedure for pulling the specimens was to use a 0.20 inch/minute rate with an extensometer that conforms to ASTM E83 class B2 with a range of at least 50%.The facture surface was evaluated using a JEOL JSM 5610 Scanning Electron Microscope (SEM) at The University of Texas at Austin with a Tungsten filament, 3nm resolution, and magnification up to 300,000 times.A total of 24 specimens were produced and evaluated. These specimens were segregated based on the tensile results, with selected samples being evaluated using SEM. Specimens were sputtered after tensile testing and prior to evaluation in the SEM.

These specimens were segregated based on the tensile results, with selected samples being evaluated using SEM. Specimens were sputtered after tensile testing and prior to evaluation in the SEM. 8 of the 24 specimens were used with redundant specimen data being discarded.Evaluation of Melt Flow Rate on Tensile Specimen Failure Using Injection MoldingIt has been observed that a primary contributor to poor mechanical properties of laser sintered components is the thermal aging, or degradation, of the material. Continuous exposure of the un-sintered material to the elevated thermal environment causes cross-linking in the polymer chain and contributes to decreased physical and mechanical properties.

Continuous exposure of the un-sintered material to the elevated thermal environment causes cross-linking in the polymer chain and contributes to decreased physical and mechanical properties. As the material undergoes thermal aging, the viscosity of the material is seen to increase and can be evaluated using a melt-flow measurement tool.A key consideration in evaluating tensile specimen failure is whether this material degradation is an intrinsic material defect or whether this change in viscosity has an effect on process specific layer-to-layer adhesion. Unusable polyamide powder was segregated from production at Harvest Technologies and its viscosity (melt flow rate) evaluated.

Unusable polyamide powder was segregated from production at Harvest Technologies and its viscosity (melt flow rate) evaluated. This used polyamide powder was then processed in an injection molding machine at The University of Texas to produce several tensile coupons.Geometric Consideration of the Interface Boundary LayerWhen laser sintered specimens are observed optically, there is a visible boundary layer between subsequently fused layers. This boundary layer is exclusively seen in the X-Y plane. This interface boundary layer appears to be the primary culprit in decreased mechanical properties and can be characterized as a layer of coplanar voids. Therefore, tensile specimens were created using CAD to simulate this boundary layer with varying degrees of severity.

Therefore, tensile specimens were created using CAD to simulate this boundary layer with varying degrees of severity. These specimens were then processed in the stronger X-direction in order to isolate the effect of the decreased z-direction mechanical properties on this experiment.Figure 3 illustrates the CAD model that was created to simulate coplanar powder particles. The tensile bar was scaled up from a standard ASTM D638 tensile bar the particle diameter was modeled at 0.100”.Figure 4: Coplanar spheres and voidsThe two halves of the tensile bar were created independently in order to allow for a variation in the interface interference between a plane on one half and coplanar spheres on the other. This interface will be defined by an “h/r Ratio” seen in Figure 5. An h/r Ratio of 0 will have the spheres just touching the plane as illustrated in Figure 4. A ratio of 1 will have no effective interface boundary and is equivalent to a solid tensile bar.

An h/r Ratio of 0 will have the spheres just touching the plane as illustrated in Figure 4. A ratio of 1 will have no effective interface boundary and is equivalent to a solid tensile bar. Eleven tensile specimens were produced with an h/r Ratio from 0 to 1 at an increment of 0.1. The material used was a polyamide 11 in an optimized production build. Optimized polyamide 11 production builds typically yield the highest mechanical properties and should provide the greatest contract for the varying h/r ratio tensile specimens.Figure 5: h/r RatioFigure 6: SEM of X Tensile Bar Fracture Surface. (35x)Figure 7: SEM of X Tensile Bar Fracture Surface. (250x)Figure 8: SEM of X Tensile Bar Top-Facing Surface. (35x)Figure 9: SEM of X Tensile Bar Top-Facing Surface. (250x)Figure 10: SEM of Z Tensile Bar - Delamination. (35x)Figure 11: SEM of Z Tensile Bar - Delamination. (250x)Figure 12: SEM of Z Tensile Bar – Brittle Fracture.

(250x)Figure 10: SEM of Z Tensile Bar - Delamination. (35x)Figure 11: SEM of Z Tensile Bar - Delamination. (250x)Figure 12: SEM of Z Tensile Bar – Brittle Fracture. (35x)Figure 13: SEM of Z Tensile Bar - Brittle Fracture. (250x)Figure 14: SEM of Y Tensile Bar - Ductile Fracture. (35x)Figure 15: SEM of Y Tensile Bar - Ductile Fracture. (250x)RESULTSThe tensile testing performed yielded the results illustrated in Table 1. A data from 8 specimens is listed in this table and illustrated in the SEM photographs. Tensile specimen nomenclature identifies the geometric location of the sample in the build as well as the build orientation. A “zx” specimen would be oriented primarily in the z- axis with the secondary axis oriented in the x-axis. An “x” specimen would be oriented flat in the x-y plane with the primary length along the x-axis.

An “x” specimen would be oriented flat in the x-y plane with the primary length along the x-axis. The x-axis specimens yield both better mechanical properties and consistent results.Table 1: Tensile Properties and Fracture ModeSampleZX 12.0.3 ZX 0.0.3 ZX 0.12.3 ZX 13.11.3 X 4.2.8 X 4.3.8 X 4.4.8Ultimate Tensile Strength (UTS) 494 psi 2969 psi 1278 psi 5164 psi 5886 psi 5991 psi 5763 psiElongation at Break (EOB)1% 1% 1% 3% 7% 6% 7%Fracture ModeDelamination Delamination Delamination Brittle Brittle Brittle Brittle+Y 1.4.01 1 Ductile fracture taken from independently optimized build.6099 psi44%DuctileMorphologyThe SEM image shown in Figure 6 shows clear patterns for the layers with elongated horizontal voids between layers. Figure 7 is a magnified region of the same sample. It illustrates a single layer with voids both above and below the fully melted layer.

Figure 7 is a magnified region of the same sample. It illustrates a single layer with voids both above and below the fully melted layer. There are also fully dense regions where there was sufficient energy to melt the layers together, creating a more homogeneous region.The surface of an XY tensile specimen is illustrated in Figures 8 and 9. This shows the upper facing surface and the surface to which the loose powder must be bonded by the laser. The average particle size is roughly 50 microns and individual particles can be seen to be melted in Figure 9, appearing like a lily pad or pancake structure. Some particles can be seen to have melted with other particles while others appear to be isolated with voids between particles. Of particular interest in Figure 9 is the 250(cid:1)m boulder looking particle in the center of the micrograph.The fracture surfaces of the Z tensile specimens are shown in Figures 10 through 15.

Of particular interest in Figure 9 is the 250(cid:1)m boulder looking particle in the center of the micrograph.The fracture surfaces of the Z tensile specimens are shown in Figures 10 through 15. Delamination is illustrated in Figures 10 and 11, brittle fracture in Figures 12 and 13, and ductile fracture in Figures 14 and 15.Injection Molded Tensile SpecimensStandard injection molding parameters from a materials handbook were used to create fully dense ASTM D638 tensile specimens. This particular machine was not in the greatest state of repair, but was sufficient in creating several tensile specimens from the used material. The powder was poured directly into the extruder and processed manually. The results of the tensile test is presented in Table 2.

The powder was poured directly into the extruder and processed manually. The results of the tensile test is presented in Table 2. Both specimens exceeded the range of the extensometer (50%) with one specimen pulled until it reached the mechanical limits of the load cell.Table 2: Injection Molded Tensile ResultsSampleSample 1 Sample 2Ultimate Tensile Strength (UTS) 5572 psi 5530 psiElongation at Break (EOB) >50% >200%Fracture Mode Ductile Ductileh/r Ratio Tensile SpecimensTwo builds of 11 samples each were built and evaluated, with one set of samples having a finished surface and the other set having no post finishing done. There was no difference in ultimate tensile strength and elongation between builds.

There was no difference in ultimate tensile strength and elongation between builds. The data presented in Table 3 and Figure 16 is from the natural tensile specimens.Table 3: h/r Ratio Tensile ResultsSample 1 2 3 4 5 6 7 8 9 10 11h/r Ratio 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0(cid:2)(cid:3) (psi) 2772 2854 3028 3056 3163 3047 2824 3090 3225 3214 3015UTS (psi) 2954 3124 4021 4976 5312 5468 5882 6130 5994 6133 7391EOB % 4 4 6 7 9 9 17 17 16 18 >50*Specimen exceeded extensometer limit of 50%.Figure 16: Stress-Strain Curve for h/r Ratio SpecimensDISCUSSION[Figure 6] With the understanding that the z-axis is the primary direction for added layers, a pattern of voids between layers is apparent. This stratification contributes to anisotropic material properties and significant weakness if a tensile stress is exerted parallel to the z-axis.

This stratification contributes to anisotropic material properties and significant weakness if a tensile stress is exerted parallel to the z-axis. In addition, these patterns of voids contribute to a significant reduction in elongation and ultimate tensile strength.[Figure 9] It appears that the 250(cid:1)m particle was formed when several particles melted together but did not fully adhere to the surface. When the particles cooled, there became a condensed grouping of several melted particles with a surrounding void. This voided region may contribute to voids in the part unless sufficient energy is used to re-melt these particles and bond with theFigure 16: Lightly adhered layers contributing to delamination.surrounding surface.It is apparent from the analysis of these samples that the cracks propagate through voids created by a lack of full melt between layers.

The general observations for each fracture mode are listed below:[Figures 10 and 11] These images show a fracture of the Z specimenFigure 17: Incomplete fusion contributing to brittle fracture.that is parallel to the x-y plane. Three distinct layers are represented as the crack propagation followed the area between two parallel layers and then migrated to neighboring layers. This pattern, whenFigure 18: Initial Layer of Sintered Powderexamined in Figure 11 shows that the particles are roughly 50 microns in diameter, which corresponds to the average particle size of the raw material. When compared to the morphology of the top layer (Figures 8 and 9), it can be seen that the bottom layer has very little bonding to the previous layer.[Figures 12 and 13] Voids are isolated and somewhat periodic, with similar size and spacing. Fracture surfaces show a ductile tear with trans-granular fractures or river marks.

Fracture surfaces show a ductile tear with trans-granular fractures or river marks. These fracture surfaces are similar to the ductile specimen (Figure 14) but with void fraction on the failure surface higher.[Figures 14 and 15] With an elongation of 44%, it may be observed that there is void nucleation where the material tore apart from itself. From Figure 15 shards of material and a very dense surface are observed.It is apparent that the material properties observed in Table 1 can be correlated to the extent of voids present between sintered layers. With very little adhesion between layers, delamination and elongation values near 1% are seen. With suffiicient energy to melt through the target layer into the previous layer, a fully dense part will yield elongations well above 10%.

With suffiicient energy to melt through the target layer into the previous layer, a fully dense part will yield elongations well above 10%. As the layer has a higher degree of melt, the voids will decrease through the brittle range until the voids no longer contribute to the fracture.Figures 3 and 4 are illustrations of two subsequently built layers with voids being created between layers through the lack of complete particle melt to the previous layer. Figure 17 shows a larger separation between layers and represents the sample seen in Figures 10 and 11 (delamination). Since the voids are coplanar, they are analogous to pre-existing cracks. The crack propagation in this example will cause a shear between layers resulting in delaminate.Figure 18 represents the sample seen in Figures 12 and 13 (brittle fracture) with voids present, but with the voids being isolated and periodic.

The crack propagation in this example is much less directional since the voids are no longer exclusively coplanar. A fracture may originate between specific layers, but will typically travel through voids in neighboring layers since the distribution of voids is much less ordered.Figure 18 is a graphical representation of an initial layer created in the laser sintering process. The depth of this layer is dependant upon the laser energy, but regardless of the amount of energy on the first layer, the downward facing surface will be defined by the raw material particle size as seen in Figure 10. The top of the layer is flat and will look like the surfaces seen in Figure 8 and 9. This irregular downward facing surface is the impetus for void creation.To insure the best properties of a part created though the laser sintering process, layers must be processed in such a way as to minimize or eliminate voids between layers.

Voided regions between layers define the fracture mode.Effect of Reused Powder on Particle MeltBased on the tensile data acquired, it was apparent that degraded powder did not yield poor mechanical properties when injection molded. This test gives us a good degree of confidence that the degradation of mechanical properties in the laser sintering process is not an intrinsic property defect but a property that affects the processing of the polyamide material in laser sintering.Coplanar Void (h/r Ratio) BehaviorThe existence of an interface boundary layer can be illustrated in Figures 6 and 7 from the SEM images. In addition, work done at the University of Louisville i with microtoming and optical microscopy of samples shows evidence of a boundary region between layers.

In addition, work done at the University of Louisville i with microtoming and optical microscopy of samples shows evidence of a boundary region between layers. The work cited here and in Figures 20 and 21 have yet to be officially published at the time of this writing.The images obtained from the University of Louisville show a cross section in a sample that is oriented the same as the fractures shown in Figure 6. The samples were created by using a microtoming technique common in medical pathology. The samples were sliced in layers roughly 5 microns thick to evaluate individual layers. It can be seen in Figures 20 and 21 that the fused layers (lighter color) are fully dense and the particles at the bottom of the layer (darker) are nylon particles that were not sintered.

It can be seen in Figures 20 and 21 that the fused layers (lighter color) are fully dense and the particles at the bottom of the layer (darker) are nylon particles that were not sintered. It can also be observed that there areFigure 20: Optical Microscopy in a Bright Field (5X)Figure 21: Optical Microscopy in a Bright Field (10XFigure 19: Trends in h/r Tensile Propertiesseveral tears in the sample shown in Figure 20 that correspond with these darker (unmelted) regions.When analyzing the graph in Figure 19, there are two distinct regions present (ductile and brittle). An h/r Ratio of 0 to 0.2 represent fracture before strain hardening which is seen in laser sintering as delamination. In addition, There is a grouping of these regions as illustrated in the stress-strain curve in Figure 16.The use of an h/r Ratio helps quantify the failure modes in a macro scale and can help in identifying material process improvements that may help on a micro scale.

This correlation is significant.CONCLUSIONSThermally degraded powder does not affect the tensile properties when used in the injection molding process. Therefore, poor viscosity powders affect the laser sintering process negatively.Poor layer-to-layer adhesion contributes to the formation of an interface boundary layer.

Therefore, poor viscosity powders affect the laser sintering process negatively.Poor layer-to-layer adhesion contributes to the formation of an interface boundary layer. This is evident in both SEM and optical microscopy.The primary cause of the lower mechanical properties in the Z-axis can be attributed to layer-to-layer adhesion.There are three distinct modes of failure in laser sintered specimens: delamination, brittle fracture, and ductile fracture.It is apparent from the analysis of these samples that the cracks propagate through voids created by a lack of full melt between layers.The use of an h/r Ratio helps qualify the regions of fracture.An h/r Ratio of 0.6 demonstrates a shift between brittle and ductile fracture modes.There is a major jump in mechanical properties when the interface boundary layer is significantly minimized between the 0.9 and 1.0 h/r Ratio region.i Timothy J. Gornet, pri.conv., University of Louisville.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEFFECT OF IN-SITU COMPACTION AND UV-CURING ON THE PERFORMANCE OF GLASS FIBER-REINFORCED POLYMER COMPOSITE CURED LAYER BY LAYERShiferaw D. Beyene*, Beshah Ayalew* and Srikanth Pilla*† *Clemson University, International Center for Automotive Research 4 Research Drive Greenville, SC 29607 †Department of Materials Science and Engineering, Clemson University, Clemson, SC, 29634AbstractIn this paper, the effect of in-situ compaction and UV-induced curing on the performance of fiber-reinforced diglycidyl ether bisphenol A (DGEBA) based epoxy is studied for a layer-by-layer curing process. The optimum percentage of photo-initiator concentration and UV-intensity were obtained by conducting a different experiment for each of them. Fourier Transform Infrared (FTIR) spectroscopy method is used to determine the degree of cure.

Fourier Transform Infrared (FTIR) spectroscopy method is used to determine the degree of cure. Then, short beam shear (SBS) test is conducted to measure the inter-laminar shear strength of the cured product under different compaction load. The UV intensity and Photo-initiator concentration were kept constant during the test. The result showed that thick composite parts fabricated with in-situ compaction and UV curing process have showed increased inter-laminar shear strength with increased compaction load up to a certain point. An increase in compaction beyond this point decreased the interlaminar-shear strength.IntroductionAdditive manufacturing (AM) of fiber-reinforced composites is currently conducted by stereolithography, layer-by-layer, fused deposition modeling, selective laser sintering, and extrusion [1].

Layer by layer also called laminated object manufacturing (LOM) manufacturing is a type of additive manufacturing process which uses sheet of material and adhesive to make thick part [2]. Due to its inherent capability for handling sheet materials such as fiber reinforced composite prepregs, for producing geometrically complex near net shape objects , LOM is preferred is preferred manufacturing process among others [3]. Therefore, in the last two decades there has been a growing interest in layer-by-layer method for composite manufacturing.The interfacial characteristics of composites fabricated by LOM was investigated by [4] and reported that interlaminar shear strength is one of the main factors affecting the overall performance of the final product. C-shaped panels, made from polymer matrix composites, were directly fabricated by curved LOM [5]. A vacuum thermoforming apparatus was applied to bond commercial prepregs.

C-shaped panels, made from polymer matrix composites, were directly fabricated by curved LOM [5]. A vacuum thermoforming apparatus was applied to bond commercial prepregs. In other research, model based UV-induced stepped-concurrent curing approach for manufacturing of thick acrylate based composites was proposed [6, 7] and later validated with experiment [8].The major issue for layer by layer manufacturing process is the difficulty of bringing adjacent layers to fully consolidate and cure [5]. It is quite important to increase the interlayer strength and reduce the void content as low as possible. A handful researchers have been studying to address1152this issue, [9] studied the effect of compaction and UV exposure on the inter-laminar shear strength (ILSS) and reported that improved ILSS of the final product is achieved when the optimum ILSS and UV intensity is chosen.

The concept of layer-by-layer in-situ UV processing concept to make a thick glass fiber reinforced polyester composite was studied by [10]. The effect of layer thickness on the inter-laminar shear strength on achieving better ILSS is studied. The LOM system used in previous studies require post processing operation to increase the bonding between layers [11].In this paper, we propose a new method of UV-induced layer-by-layer composite manufacturing which uses in-situ curing and compaction to increase the ILSS. First, the optimal curing time sequence that is the switching, layering and final times are determined by optimal time control. Using these optimal times, four different samples were prepared with different pressure applied during the curing. The final degree of cure was checked by observing the FTIR of the cured samples.

Using these optimal times, four different samples were prepared with different pressure applied during the curing. The final degree of cure was checked by observing the FTIR of the cured samples. The inter-laminar shear strength of these samples are then investigated by short beam shear test (SBS).The rest of the paper is organized as follows: Section II discusses the UV curing process model, Section III explain how the optimal switching time control is done. Section IV discusses the experimentation. Section V discusses the results and conclusions are included in Section VI.Cure process modelFor two-layer curing, a one-dimensional heat conduction for UV-induced cationic curing process can be illustrated by the schematic of Fig.1.

Where, the first mode is from initial time till UV off, the second mode is from UV off time till the second layer is added; the third mode is from layer addition (UV on) time till the UV off time on two-layer curing and final mode is from UV off time till final cure time. At the top surface, a convective boundary condition is taken (BC1) and, insulation boundary condition is considered at the bottom and sides (BC2). As depicted in the figure, there are two switching modes: switching mode one (MS1) which is from UV on to UV off and switching mode two (MS2) which is when a layer is added on the exiting layer.Figure 1. Schematic of cationic curing process for two layers1153The process model for UV-induced cationic polymerization consists of the cure kinetic model for curing under UV and curing without UV (UV is off)). The model consists of the exothermic reaction (heat generation), conduction and convention heat transfer.

H∆ is enthalpy of polymerization; is degree of cure at time t and depth y ; l is the thickness (depth) of the sample; A is pre-exponential constant; E is the activation energy ; R is universal gas constant; absT the absolute temperature in Kelvin; 0I is the initial UV intensity; λ is ft are switching and final cure times. UV attenuation constant;( ,)t yα ( ,)ik is initiation rate constant; st andThe physical parameters used for simulation are summarized in Table 1.TABLE1: Physical parameters for the numerical simulationFraction by volumeρ[g/cm3]c [J/(kg K)]yk [W/(cm oc)]H∆ [J/g]Resin0.41.151.750.0025332Glass fiber0.62.5750.800.01275Composite121.020.0037332UV TransmissionThe UV transmission of the epoxy resin is found by comparing the UV intensity at the top surface with the one measured at the bottom of the sample. The transmission for different photo-1154initiator (PI) concentration is shown in Fig. 2.

The transmission for different photo-1154initiator (PI) concentration is shown in Fig. 2. As seen in the figure the maximum absorbance is at 4% however, since the difference in the absorbance for 3% PI and 4% PI are not significantly different (<1mW/cm2). Moreover, studies have shown that increasing the PI beyond certain level will not benefit the final degree of cure [17]. Therefore, for this study we opt for 3 % which is in agreement with literatures [18].Figure 2. UV transmission for different PI concentrationOptimal switching time controlAn optimal control defined as one that minimizes/maximizes some performance criteria [19]. In this paper, the objective of optimal control is to get a uniform final cure distribution. The performance measure is therefore related to what extent the actual final cure distribution deviates from the desired cure distribution. The performance measure is then given by Eq.

The performance measure is therefore related to what extent the actual final cure distribution deviates from the desired cure distribution. The performance measure is then given by Eq. (5)J=1 2l∫0( α ( tf)−αdes)2dywhere J is the performance index;)ftα is the vector containing final degree of cure at all points desα is the desired degree of cure along the depth at the final time(considered along the depth and ft. l is the thickness (depth) of the sample.Two fiber-reinforced DGEBA based prepregs with each layer 2mm thickness were considered for this study. The process input, UV intensity, is kept uniform for the first mode and turned off at the first switching time (MS1) and then it is left to cure by itself till the second switching time (MS2) this process continues for any number of layers. A one-dimensional UV-induced cationic polymerization model given in Eq. (1-4) is used to get the temperature and cure distributions along the depth.

A one-dimensional UV-induced cationic polymerization model given in Eq. (1-4) is used to get the temperature and cure distributions along the depth. The switching times and final time are the control variables that can be optimized to get a minimum deviation of the final cure from the desired value (98%).In our previous study, we derived the necessary conditions for optimality [16]. Similar derivations can be found in [6, 7]. The necessary conditions for optimality and the coupled PDE- ODE equations (1-4) are taken as constraints for the performance index given in Eq. (1). Steepest descent algorithm [19] was applied to find the optimal switching times and final time. The optimal switching and final times found from the optimal time control are given in Table 2.1155[5]TABLE 2.

The optimal switching and final times found from the optimal time control are given in Table 2.1155[5]TABLE 2. Optimal switching and final times.Mode Optimal switching time (s)1 2882 5053 8004 884ExperimentationIn-situ curing and compactionComposites were fabricated by irradiating the prepreg samples using a 16.1W Clearstone UV LED which has UV emission peak wavelength of 365nm. The amount of UV reaching to the surface of the resin sample is measured by using a digital UV radiometer (Solarmeter) whose resolution is 0.1 mW/cm2. All samples were placed at the same depth from the UV LED. In our previous work [15], we verified experimentally that the optimal switching times found from simulation help to get through cure for 2mm thick resin. Hence, in this study 2mm thickness is used for a single layer. The schematic of in-situ curing and compaction is shown in Fig. 3.

Hence, in this study 2mm thickness is used for a single layer. The schematic of in-situ curing and compaction is shown in Fig. 3. As depicted in the figure, the first layer is cured and then a new layer is added with a clear glass sheet on it where four equal loads are applied outside the UV exposure area. The two layers are then cured and consolidated with the in-situ pressure applied on the clear glass sheet.Figure 3. Schematic of in-situ compaction and curingFinal degree of cureTo make sure that the samples are completely cured, we recorded the FTIR of all cured samples and compared with the FTIR of the uncured resin. The characteristic absorptions of oxirane ring at 914.9 cm-1 is observed in the epoxy resin (before curing). This is attributed to the C-O deformation of oxirane group [20, 21]. As seen in Fig.

The characteristic absorptions of oxirane ring at 914.9 cm-1 is observed in the epoxy resin (before curing). This is attributed to the C-O deformation of oxirane group [20, 21]. As seen in Fig. 4 the disappearance of the band in all samples, which were manufactured with different in-situ pressure, shows that all samples are completely cured [21].1156Figure 4. FTIR of all samples comared with uncured epoxy resinEffect of pressure on ILSSTo study the effect of compaction on the interlaminar shear strength, four samples were fabricated by the in-situ curing method with varying pressure applied to the samples. Then short beam shear (SBS) test was used to measure the interlaminar shear strength for each case. The ILSS is calculate as follows [22]:ILSS=0.75×mP × b hwhere ILSS is short-beam strength in MPa; Pm is the maximum load recorded during the test in Newton (N); b is the measured specimen width in mm; and h is the measured specimen thickness in mm.

The maximum observed load during the SBS test and the corresponding ILSS are summarized in Table 3.Table 3: Maximum observed load and the corresponding ILSSSample no. Applied load (N) Maximum observed load (N)1 2 3 40 6.8 18.2 33.9712.8 779.8 916.5 864.5ILSS (MPa) 16.7 18.3 21.5 20.3Figure 5 a) shows the maximum load observed for different loading conditions and Fig.5 b) shows the correlation between ILSS and applied load during in-situ curing. The ILSS increased by 38.75 % when the applied load increased from 0 to 18.2 N and decreases when a sample is cured under higher load. This shows that an optimum pressure is required to get better interlaminar shear strength and for our case.1157[6]1050241000950a22b900N ( d a od e v r e s b o m u m x a Mi850800750) a P M S S L I(2018700160102030400102030Applied load (N)Applied load (N)Figure 5.

Maximum observed load (a) and ILLS (b) for different loading conditions.ConclusionIn this paper, the effect of in-situ compaction and UV-induced curing on the performance of fiber-reinforced DGEBA based epoxy is studied for a layer-by-layer curing process. The optimum percentage of photo-initiator concentration is chosen for a given UV-intensity by conducting different experiment. Fourier Transform Infrared (FTIR) spectroscopy method was used to check the completeness of the final degree of cure. Then, short beam shear (SBS) test was conducted to measure the inter-laminar shear strength of the cured product under different compaction load. The curing times were found from optimal switching time control and the UV intensity and photo- initiator concentration were similar for all samples.

The curing times were found from optimal switching time control and the UV intensity and photo- initiator concentration were similar for all samples. The result showed that thick composite parts fabricated with in-situ compaction and UV curing process have showed increased inter-laminar shear strength with increased compaction load up to a certain point. An increase in compaction beyond this point decreased the interlaminar-shear strength.AcknowledgmentThis research has been supported by the National Science Foundation under grant number:CMMI-1537756References1. Kruth, J. P., Levy, G., Klocke, F., & Childs, T. H. C. (2007). Consolidation phenomena in laser and powder-bed based layered manufacturing. CIRP annals, 56(2), 730-759.2. Wong, K. V., & Hernandez, A. (2012). A review of additive manufacturing. ISRN Mechanical Engineering, 2012.3. Mailunas, R. (2001). Laminated Object Manufacturing-Based Design Ceramic Matrix Composites. NORTHROP GRUMMAN CORP BETHPAGE NY.4.

This effect, influenced by the bond strength in the z-direction, can be more pronounced for components with longer layer times, as the bottom layers tend cool below the glass transition temperature (Tg) of the material, thereby restricting thermal fusion between the printed layers. The work discussed here builds on the previous work by the authors, demonstrating infrared preheating as a technique to actively control the bond temperature during printing to improve the mechanical properties (z-direction) of parts printed on a large-format extrusion AM system. IR preheating was used on the surface of printed layers just prior to the deposition of the next layer to increase the surface temperature closer to the glass transition temperature. The current study explores the effect of variation in bead surface temperatures (indicative of variations in layer times) prior to and after pre-heating on the mechanical properties in z-direction.1.

INTRODUCTION1.1 MotivationExtrusion deposition additive manufacturing, the process where the feed material is melted and deposited layer-by-layer on a build platform, is now being used to print components using a variety of polymers and composite materials on print platforms with build volumes ranging from ~ 10 cm3 up to ~130 m3 [1, 2]. Despite several advancements in the development of new materials and printers based on this process, one of the challenges that still exists is the inability to obtain printed components with isotropic mechanical properties [3-5]. There exists a certain level ofThis manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05- 00OR22725 with the U.S. Department of Energy.

There exists a certain level ofThis manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05- 00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, worldwide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes.

The Department of Energy will provide public access to these results of the DOE Public Access Plan federally (http://energy.gov/downloads/doe-publicaccess-plan).sponsoredresearchin accordance with1576anisotropy in the properties of printed parts (mechanical properties being lower in the transverse direction or the print direction when compared to the axial direction) as the layer-wise deposition process makes these properties dependent on temperature history and processing conditions [6-8]. During printing on a large format AM platform (6 m x 2.4 m x 1.8m) [1], without an active temperature control of the entire build volume, increasing part geometry (or layer times) would lead to increased cooling of the deposited bottom layers (or beads) before the next layer gets deposited.

If the bonding mechanism between the deposited polymer layers is primarily thermal fusion and polymer inter-diffusion, then the temperature of the bottom layer (or substrate) should at least be above the glass transition temperature (Tg) of the polymer for a sufficient time period to enable good thermal bonding [6, 7]. However, for large-scale system, this condition in not always fulfilled and therefore, there is a need to actively control the bond temperature during printing to obtain better mechanical properties in the print direction.1.2 Infrared PreheatingAs a strategy to address this issue of cooling of bottom beads below Tg before the next bead gets deposited for large-scale AM systems, the authors previously investigated and reported infrared (IR) preheating technique to be an effective method to improve transverse direction mechanical properties (z-direction) [9, 10].

An experimental design feasible for using IR preheating on large-scale printers such as the Big Area Additive Manufacturing system was demonstrated, and this technique was found to be a useful to modify the interlayer bond temperature actively during printing, independent of the material used and part geometry. The previous test set-up comprised of infrared lamps, pyrometers and the extruder all designed to move as a single unit while printing. The lamps were fitted such that they preceded the extruder to heat up the bottom layers just prior to the deposition of the next layer. Two pyrometers, one positioned before the lamp and another after, were used to record the cold bead (substrate) temperature prior to heating (Tcold) and the substrate temperature after IR preheating (Thot), just prior to deposition of next layer, respectively.

The previously performed mechanical tests results overall indicated the IR preheating technique to be beneficial for print conditions where the cold substrate temperature (Tcold) was reduced below Tg (in the range of 75-85 oC and Tg of the material printed ~ 110 oC) and when IR preheating increased the substrate temperature closer to or slightly above Tg.Despite the reported efforts so far by the authors on using the IR preheating technique to actively control the bond temperature, there was a need to refine the experimental set-up and print conditions to have a better control of the substrate temperature before and after preheating to be able to better understand the effect of preheating on the mechanical properties. The previous efforts primarily used variations in print speed, intensity of the lamps (heat input), and extrusion temperature as methods to modify substrate temperature before and after preheating.

The previous efforts primarily used variations in print speed, intensity of the lamps (heat input), and extrusion temperature as methods to modify substrate temperature before and after preheating. There was no other active control of substrate temperature before preheating, and there were several cases of excessive heating caused by the lamps which might have led to some possible degradation of the polymer, thereby in turn weakening the z-direction properties. To have a better understanding of the conditions where this technique could be very effective, in the work reported here, print trails were performed by actively controlling the substrate temperature before and after preheating.

Printing with a set speed and deposition temperature, artificial cooling was used to create different cold substrate temperatures (by varying the number of cooling cycles) and the intensity of the lamp was constantly controlled to raise the substrate temperature to a certain set value (above Tg) by preheating. This reduced overheating issues as well as provided a way to alter the cold substrate1577temperature to ranges well below Tg (40 to 80 oC) unlike the previously reported trials. Also, it should be noted that lower cold substrate temperature is indicative of larger parts or increased layer times, thereby making these studies more useful to understand the effect of preheating in case the of printing very large parts.2.

EXPERIMENTATION2.1 Experimental Set-upFigure 1 represents the modified experimental set-up on the large-scale AM system used in this work comprising the IR lamps, pyrometers, extruder, all inter-connected and moving as a single unit, as well as provisions for artificial cooling during printing.Figure 1. Experimental set-up indicating (a) various key components and (b) IR preheating during printing.The feed material used for printing was the same as the previous trials, acrylonitrile butadiene styrene (ABS) reinforced with 20 wt.% short carbon fiber. The IR lamp used was a Strip IR model number 5306B-02-1000-01-00 fitted at a standoff distance of 10 mm from the bead surface. Components printed under different conditions were 0.25 m high hexagons with dimensions as shown in Figure 2. Three hexagons were printed, with cold substrate (Tcold) target temperatures (i.e.

Components printed under different conditions were 0.25 m high hexagons with dimensions as shown in Figure 2. Three hexagons were printed, with cold substrate (Tcold) target temperatures (i.e. substrate temperature just prior to preheating) being 40 oC, 60 oC and 80 oC controlled by varying the number of active cooling cycles. Active cooling was achieved through forced convection. Eight sets of insulated hoses were used to direct air flow over the printed structure. The target substrate temperature after IR preheating was set to 150 oC (above Tg) for all the cases. The print speed for all the parts was set to 5.1 cm/s.1578Dired:ion of extruder motion1.17IR preheated sideLampsonLamps offNon-heated sideFigure 2. Schematic of the printed hexagonal sample (all dimensions in m), sample height: 0.25 m.2.2 Tensile TestingTensile testing was performed by harvesting samples from both, heated and the non-heated sides (control samples) of the hexagon.

Plaques were cut out from the sections harvested and tensile test specimens were milled from these plaques. The target specimen dimensions are shown in Figure 3 and were in accordance with ASTM D638 (Type I). Tensile tests (z-direction of the sample) were performed using an MTS 4.4 kN (1 kip) load cell and MTS 200k micro-stain extensometer at a testing speed of 1.524 mm/min (0.06 in/min) until failure...57,, M ..... ..lI◄115165►IFigure 3. Tensile test specimen dimensions (mm), sample thickness: 6.35 mm.3. RESULTS AND DISCUSSIONTable 1 summarizes the printing and preheating conditions varied in the previous reported print trials to modify the substrate temperature before and after preheating (Tcold and Thot) and Figure 4 summarizes the tensile test results reported for preheating conditions that used a 1kW IR lamp with variations in print speeds and the extrusion temperature [9,10].

In almost all the reported cases, there were no statistically significant improvement in tensile strength due to preheating. In addition, for many conditions, IR preheating lowered the average z-direction tensile strength to some extent which was attributed to the possibility of excessive heating (as the substrate was preheated up to the chosen extrusion temperature) leading to possible degradation of ABS in the process.1579The tensile test results for the samples harvested in the present work, with a good control of Thot and Tcold, are as shown in Figure 5. The data clearly indicates an increase in the average tensile strength for all the three Tcold conditions, with the increase being statistically significant for the case with Tcold of 40 oC. The average z-direction tensile strength for this condition was increased by 81% due to IR preheating.

The average z-direction tensile strength for this condition was increased by 81% due to IR preheating. These trends show that IR preheating is a very useful technique for improving bond strength, especially for print scenarios involving long layer times that cool the part down to temperatures significantly below the glass transition temperature of the material.Table 1. Process parameters varied in previous trials.Print speed (cm/s) Extrusion temperature (oC) 215, 235, 250 Lamp power setting3.8, 5.1, 7.680-90% intensity for 1kW lamp 2 X 500 W lamps (at max intensity) 8, 2.5, 1Standoff distance (cm)20Non-heated side IR preheated side'ii' 15 IIL:::E -.c -m ! -(I)C 10~ 4ft C ~50Print speed (an/s} 3.85.17.63.85.17.63.85.17.6Texbusion: 235 °CTemusion: 250 °CFigure 4. Summary of tensile test results from previous trials [10].158012T hot= 150 °C (after preheating)_ 10 Cl:I ll.

Summary of tensile test results from previous trials [10].158012T hot= 150 °C (after preheating)_ 10 Cl:I ll. ~ 8J: -C) f -U)C:6Non-heated side IR preheated side~"' C: ~420Tcold = 40 ocTcold = 60 °CTcold = 80 ocFigure 5. Ultimate tensile strength (z-direction) for the three cooling conditions.4. CONCLUSIONSPrevious work by the authors demonstrated infrared preheating to be a useful technique to enhance the z-direction mechanical properties of components printed on large-format AM platforms. However, in those print trials, the substrate temperature (bottom beads) before and after preheating was controlled only by varying the print parameters and heat input conditions. The cold bead temperatures obtained by varying these parameters, although below the Tg of the material, were in the range of 75 – 85 oC and excessive preheating in some cases led to the possibility of polymer degradation.

In this work, with modifications in the experimental set-up and processing conditions, the authors were able to simulate longer layer times (or larger part geometry) by actively cooling the substrate to chosen temperatures significantly below Tg and then preheating up to a certain temperature well above Tg (also below the extrusion temperature). These conditions enabled better control over the process as well as led to the possibility of printing parts where IR preheating enhanced the tensile strength by even ~ 80 %. Increase in the average tensile strength was observed to be greater as the substrate temperature lowered. This therefore indicates that the IR preheating technique to be advantageous for large-scale prints involving longer layer times. Future work involves experimental trials with multiple preheating temperatures (Thot) as well as validation of the trends in properties by performing mechanical testing using other methods.5.

Future work involves experimental trials with multiple preheating temperatures (Thot) as well as validation of the trends in properties by performing mechanical testing using other methods.5. ACKNOWLEDGEMENTSThe authors gratefully acknowledge Dr. Donald Erdman at Oak Ridge National Laboratory for the use of lab facilities for mechanical testing and Techmer ES for providing the materials used in this work. Research was sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle, LLC.15811.2.3.4.5.6.7.8.9.10.6. REFERENCESDuty, C.E., Ajinjeru, C., Kishore, V., Compton, B., Hmeidat, N., Chen, X., Liu, P., Hassen, A.A., Lindahl, J. & Kunc, V. "What makes a material printable? A viscoelastic model for extrusion-based 3D printing of polymers." Journal of Manufacturing Processes 35 (2018): 526-537. https://doi.org/10.1016/j.jmapro.2018.08.008.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEFFECT OF INTER-LAYER COOLING TIME ON DISTORTION AND MECHANICAL PROPERTIES IN METAL ADDITIVE MANUFACTURINGY.K. Bandari*, Y.S. Lee, P. Nandwana, B.S. Richardson, A.I. Adediran, L.J. LoveManufacturing Demonstration Facility, Oak Ridge National Laboratory, Knoxville, TN, 37932, USAAbstractLaser Metal Deposition with wire (LMD-w) is one of the novel Direct Energy Deposition (DED) processes that is gaining the attention of various industries, especially aerospace, due to the potential cost and lead time reductions for complex parts. However, subjects of development include optimization of process parameters (for example laser power, wire feed speed, robotic travel speed, inter-layer cooling time etc.) for large scale adaption of this process.

These parameters influence residual stress which potentially results in distortion and subsequent mechanical properties. Inter-layer cooling time is one of the main influences on production volume and is typically used to help control the cooling conditions to mitigate part distortion. Therefore, this paper is aimed at investigating different inter-layer cooling times on distortion and resulting mechanical properties of the parts produced by LMD-w. Distortion of deposited Ti-6Al-4V walls was measured automatically using a laser scanner, which was attached to the robotic arm itself. Finally, suitable recommendations are discussed to optimize the inter-layer cooling time to produce parts with desired mechanical properties.Keywords: Additive Layered Manufacturing, 3-D Printing, Direct Energy Deposition, Laser Metal Deposition with wire, Ti-6Al-4V, Laser Scanner, Distortion, Residual Stress1.

IntroductionSince their first introduction more than 30 years prior, additive manufacturing (AM) processes for metals have been produced and commercialized in the market under different names, for example, Direct Metal Deposition [1], Laser Engineered Net Shaping [2], Shaped Metal Deposition [3], Selective Laser Melting [4], and Electron Beam Freeform Fabrication [5]. In every one of these processes, a heat source is used to create a melt pool into which a wire or powderized feedstock is fed, or the heat source is applied to a powder bed to create beads after solidification. The beads are realized by relative movement of the melt pool and the substrate by using an industrial robotic arm or a gantry system [6]. A part is then manufactured by depositing beads adjacent to each other and layer upon layer. The most common technique to date has been to use a high-power laser as the heat source with metal powder as the feed stock.

The most common technique to date has been to use a high-power laser as the heat source with metal powder as the feed stock. Other conventional welding techniques have likewise been introduced, for example, Tungsten Inert Gas (TIG) welding [7], Gas Metal Arc Welding (GMAW) [8], Plasma Transferred Arc (PTA) welding [9], [10], and Electron Beam (EB) welding [11].The high-power laser-based technique remains the most common because of several advantages, for example, the relatively low heat input to the base material and good shape accuracy of the realized geometries. This Laser Metal Deposition (LMD) process has been dominated byNotice of Copyright. This manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy.

This manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes. The Department of Energy will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).425powder-based technologies because of flexibility and robustness of the process. However, the powder-based technologies have been developed for building small and complex geometries, with very little focus on deposition rate.

However, the powder-based technologies have been developed for building small and complex geometries, with very little focus on deposition rate. The deposition efficiency varies depending on the system set- up, and post treatment of the scattered powder is required. For large scale parts with moderate complexity, such as flanges or bosses, it is more rewarding to use wire-based processes, since these lead to higher deposition rates, better material quality and better surface finish [12]. While the utilization of powder requires specially planned and designed feeding nozzles, the wire-based processes require much simpler equipment and standard welding hardware.Laser Metal Deposition with wire (LMD-w) is a novel process that fabricates components using a high-power laser source and adds material in the form of metal wire. The laser generates a melt pool on the substrate material into which the metal wire is fed and melted, forming a metallurgical bound with the substrate [13].

The laser generates a melt pool on the substrate material into which the metal wire is fed and melted, forming a metallurgical bound with the substrate [13]. By moving the laser processing head and the wire feeder, i.e. the welding tool, relative to the substrate a weld bead is formed during solidification. The relative movement of the welding tool and the substrate is made using a 6-axis industrial robot arm. The formation of a weld bead is illustrated in Figure 1 along with images from the real process.Figure 1 Left: Illustration of laser-wire interaction Right: Top and side view images of the real process [13]One main consideration for the resulting parts of the LMD-w process is the dimensional precision compared to an initial input geometry. Each layer is formed by melting individual passes from the wire, which experiences rapid heating, melting, solidification, and cooling during the deposition process.

Each layer is formed by melting individual passes from the wire, which experiences rapid heating, melting, solidification, and cooling during the deposition process. While the part is manufactured, the deposited material experiences continuous heating and cooling cycles as more passes and layers are deposited. One of the consequences of the thermal gradients induced in the components by the layer-by-layer deposition of material in AM processes is the build-up of undesirable levels of residual stress and distortion [14].4262. Literature ReviewThe development of residual stress and distortion in AM processes has several similarities with multi-pass welding.

Literature ReviewThe development of residual stress and distortion in AM processes has several similarities with multi-pass welding. Several researchers have explored techniques to reduce distortion in similar multi-pass welding processes including the investigations by Michaleris and DeBiccari [15] who demonstrated that decreasing heat input would result in a decrease in workpiece distortion and Masubuchi [16] who confirmed that constraining the workpiece can limit distortion. Futhermore, Deo and Michaleris [17] examined the impact of heating the weld region immediately prior to welding and found that the technique can be utilized to achieve zero net distortion. In AM research Klingbeil et al. [18] and Cornin et al. [19] each found that bulk substrate preheating can be used to decrease distortion in deposited workpieces.In AM processes, the influence of change in path planning and inter-layer cooling time between passes on residual stress and distortion have also been studied.

For instance, the selection of different deposition paths has been shown by Nickel et al. [20] to notably affect distortion in AISI 1117 C-Mn steel, and has also been studied by Mercelis and Kruth (2001) to influence residual stress in AISI 316 austenitic stainless steel. With respect to changing inter-layer cooling time between passes, Jendrzejewski et al. (2007) reported that shorter delays reduce the measured residual stress in deposited Cobased stellite SF6 alloy, and Klingbeil et al. [18] concluded delays to reduce distortion in AISI 308 austenitic stainless steel. Fessler et al. [21] reported that allowing deposition to cool between passes decreases distortion of nickel-base INVAR alloys and austenitic stainless steels. Costa et al. [22] and Torries et al. [23] studied the effect of introducing inter-layer cooling time between deposited layers on the resulting thermal history and microstructure of laser deposited AISI 420 steel and Ti-6AL-4V respectively.

The study concluded that low inter-layer cooling times result in higher temperature levels and notably altered microstructure when compared with longer inter-layer cooling times.Whereas these studies have concentrated on post-process measurements of accumulated distortion, others have used in-situ measurements to monitor the temporal accumulation of distortion. Lundbäck [24] estimated in-situ distortion on single wall builds utilizing an optical measurement system to validate their model of the GTAW process. Plati et al. [25] utilized a linear variable differential transformer to measure the deflection of the free end of a cantilevered substrate during powder-based cladding. Grum et al. [26] recorded in-process strain using resistance measuring rosettes to measure strain accumulation during laser cladding. Ocelik et al.

Grum et al. [26] recorded in-process strain using resistance measuring rosettes to measure strain accumulation during laser cladding. Ocelik et al. [27] utilized digital image correlation to measure distortion of single and multi-bead Nanosteel, Eutroloy 16012, and MicroMelt 23 powder laser depositions on C45 steel and 301 stainless steel substrates. The in-situ measurements taken in these works offer more prominent understanding into distortion accumulation than conventional post-process distortion measurements, but do not compare the impact of changing materials or inter-layer cooling time. The in-situ distortion measurements have primarily been conducted for model validation. Also, some of the other researchers Ding et al. [28], [29], [30], Montevecchi et al. [31], Simunovic et al. [32] have developed thermo-mechanical models to predict residual stress and the resultant distortion in the large parts.

[28], [29], [30], Montevecchi et al. [31], Simunovic et al. [32] have developed thermo-mechanical models to predict residual stress and the resultant distortion in the large parts. However, the relationship between inter-layer cooling time and heat input was not discussed and then interlayer cooling time is one of the parameters which affects the overall production rate.Much of this previous work on the accumulation of distortion during the deposition has focused on welding and laser cladding processes. AM, on the other hand, involves the deposition of427multiple layers and potentially large volumes of material to produce discrete shapes and components. In cases where AM processes have been investigated, the impact of changes in the inter-layer cooling time between multiple passes on the accumulation of distortion and the impact on the mechanical properties of the realized components has not been investigated.

In the work reported here, in-situ measurements of distortion during the deposition of Ti-6Al-4V are made to investigate the effect of inter-layer cooling time on the mechanical properties.3. Experimental ProcedureThis section describes the base materials, consumables, deposition conditions, power supplies, set-up and techniques deployed during the investigation. The experiments were performed using Ti-6Al-4V plates and delivered according to the specifications AMS 4911; a 1.6 mm diameter Ti- 6Al-4V solid wire consumable was used and delivered according to the specifications AMS 4954.

The compositions in wt.% for substrate and electrode wire are shown in the Table 1.Table 1 Chemical composition of substrate and wireElement Carbon Iron Aluminum Titanium Vanadium Hydrogen (ppm) Nitrogen OxygenSubstrate (wt.%) 0.01 0.16 6.35 Remainder 4.13 25 0.01 0.19Wire (wt.%) 0.02 0.18 6.03 Remainder 4.08 25 0.01 0.183.1 Preparation of substrateThe titanium substrates were cut from a large titanium plate with a thickness of 6.4 mm. The dimensions of the substrate used for the investigation was 500 mm x 155 mm. The substrate’s surface was first wiped using an acetone dampened clean paper cloth. Then a finishing disk Cibo SAG/5/115 was mounted on a hand-held grinder with a low rotational speed and the surface was ground to eliminate oxide layers and impurities. Finally, degreasing was repeated using the acetone dampened clean paper cloth. The substrate was clamped onto a steel block with ceramic washers in between them to avoid heat transfer through conduction.

The substrate was clamped onto a steel block with ceramic washers in between them to avoid heat transfer through conduction. The steel block was then placed onto an aluminum table. A schematic diagram of the arrangement is shown in Figure 2.3.2 Equipment set-upAll the investigations were conducted with a 6-axis KUKA ‘KR90 2900 extra HA’ robot. Two fiber delivered diode lasers each with the power of 10.7 kW and wavelengths of 940 – 980 nm and 1020 – 1060 nm were used, and they were combined using a laser head combiner to form a laser with 20 kW power. The laser light was conveyed to the optics to eventually form a 7.69 mm diameter out of focus laser spot on the work piece. Abicor Binzel was used as the wire feeder.428All the investigations were performed in a closed argon environment with less than 250 ppm O2 content.

Abicor Binzel was used as the wire feeder.428All the investigations were performed in a closed argon environment with less than 250 ppm O2 content. The equipment set up is shown in Figure 3.A laser line scanner was added to the system to measure the height of deposited layers, and it enabled the implementation of layer height control. It scans from the left end of the plate to the right end of the plate and records data for every 0.5 mm of travel. The laser scanner then generates layer height maps; subsequently, the correction data for the subsequent layer is generated from the control algorithm generated in-house, to maintain a set layer height for all the layers. The control algorithm was developed using LabVIEWTM software and was executed in a computer.

The control algorithm was developed using LabVIEWTM software and was executed in a computer. LabVIEWTM software was also used to interface with the robot, laser and wire feed systems as well as collecting and logging data.(a) Top view(b) Side view Figure 2 Arrangement (a) Top view (b) Side view4293.3 Design of ExperimentsThe weld parameters used for deposition are listed below in Table 2. The dimensions of the walls were 200 x 20 x 11 mm3. Continuous weld beads 200 mm long were deposited one layer at a time in the same deposition direction. With the weld parameters listed below, the single bead width and single bead height ranged between 9-12 mm and 0.8-1.2 mm respectively.

With the weld parameters listed below, the single bead width and single bead height ranged between 9-12 mm and 0.8-1.2 mm respectively. Thus, two beads were deposited next to each other with an overlap that resulted in a flat deposited layer of width 21 mm.Figure 3 LMD-w cellTable 2 Deposition and weld bead parametersParameter Laser power Robot travel speed Deposition rate Single bead width Single bead heightValue 11.5 kW 12 – 14 mm/s 2.0 +/- 1.0 kg/hr 9-12 mm 0.8-1.2 mmBefore depositing a new layer, the previously deposited layer was scanned using the laser line scanner, and the height data was recorded. To maintain a constant layer height, the control algorithm created new weld parameters which were then used to deposit the subsequent layer automatically. In total six walls were deposited. Three of them were clamped all-round the substrate, and three were clamped only at the center.

In total six walls were deposited. Three of them were clamped all-round the substrate, and three were clamped only at the center. Interlayer cooling time is the wait time between the end of depositing a layer and start of height scan for that layer. Each wall was deposited using a different interlayer cooling time. In total ten layers were deposited for each wall. The six walls are denoted as Wall 1, Wall 2 and so on until Wall 6 and are described below.430Wall 1 was built with clamps all-round the substrate with no inter-layer cooling time. Wall 2 was built with clamps all-round the substrate with an inter-layer cooling of 30 min after depositing every five layers. Wall 3 was built with a center clamp with no inter-layer cooling time. Wall 4 was built with clamps all-round the substrate with an inter-layer cooling time of 30 min after depositing every two layers. Wall 5 was built with center clamp with an inter-layer cooling time of 30 min after depositing every two layers.

Wall 5 was built with center clamp with an inter-layer cooling time of 30 min after depositing every two layers. Wall 6 was built with center clamp with an inter-layer cooling time of 30 min after depositing every five layers. These are described in the form of a table below (Table 3).Table 3 Description of walls depositedSampleWall 1 Wall 2 Wall 3 Wall 4 Wall 5 Wal 6Clamped/unclampedClamped Clamped Unclamped Clamped Unclamped UnclampedInter-layer cooling time of 30 min after every ‘n’ layers of deposition (n) 10 5 10 2 2 54. Results and discussionBelow are the graphs that show a correlation between the layer number and vertical deformation for each wall.

Results and discussionBelow are the graphs that show a correlation between the layer number and vertical deformation for each wall. Figure 4 and Figure 5 show the vertical deformations of the plate at the left end and right end of the plate respectively with respect to the layer number.Figure 4 Deformation of left end of the plate with an increase in layer number431Figure 5 Deformation of right end of the plate with an increase in layer numberThe vertical deformation on layer zero indicates the deformation of the plate without any deposited material. Likewise, the vertical deformation on layer 10 indicates the deformation of the plate after 10 layers of deposition and 30 min of cooling time.As shown in Figure 4 and Figure 5, the walls built using only center clamps experienced more vertical deformation on the right end of the plate than on the left end of the plate; the vertical deformation also increased as more layers were deposited. This is due to the preheating effect.

This is due to the preheating effect. The robot moves from the left end to the right end of the plate depositing beads from left to the right. Therefore, the heat also transfers from left to right and downwards through conduction mode primarily. There is more heat accumulation on the right end compared to the left end, which is also seen from the thermocouple readings, that were attached on the both ends of the plate. Subsequently, there is more residual stresses and thus, more deformation. Therefore, the deformation at the right end of the plate is more than the deformation at the left end of the plate.It is also seen that deformation increases with both an increase in layers and the introduction of inter-layer cooling time; the, deformation decreases after depositing the subsequent layer. From Figure 6, it is seen in the case of Wall 5, which was built using only center clamps and with an inter-layer cooling time after every two layers.

From Figure 6, it is seen in the case of Wall 5, which was built using only center clamps and with an inter-layer cooling time after every two layers. The deformation before the third layer of deposition was 4.5 mm, and then it decreased to 3.5 mm just before fourth layer of deposition. A similar trend was seen for Wall 3 and Wall 6, too. This trend is due to the shrinkage of the deposited metal, which causes the plate edges to lift resulting in vertical deformation.

A similar trend was seen for Wall 3 and Wall 6, too. This trend is due to the shrinkage of the deposited metal, which causes the plate edges to lift resulting in vertical deformation. Then when a subsequent layer is deposited after the inter-layer cooling time, the deposited metal adds heat to the previously deposited layers, which then expand and lower the edges of the plate [14].432In case of walls built with clamps all-round the substrate, the deformation was found to be nearly zero, both at the left and right ends of the plates irrespective of inter-layer cooling times.After depositing 10 layers with inter-layer cooling time and allowing them to cool down, all the Ti-6Al-4V plates were un-clamped. Then, the distortion at the right end of the plate was measured. Figure 6 shows the distorted plates after un-clamping.

Then, the distortion at the right end of the plate was measured. Figure 6 shows the distorted plates after un-clamping. The vertical deformation at the right end of the plate for each wall is measured using a ruler and is presented in Table 3.Wall 1Wall 2Wall 3Wall 4Wall 5Wall 6Figure 6 Ti-6Al-4V Walls after unclampingTable 4 Vertical deformation of un-clamped deposited wallsSample ID Wall 1 Wall 2 Wall 3 Wall 4 Wall 5 Wall 6Vertical deformation (mm) 6 7 8 7 12 9All six deposited walls were used to extract samples for room temperature tensile testing in accordance with the ASTM E-08 13a standard. One half of the wall was tested in an as-deposited condition, while the other half was subjected to the heat treatment recommended by the AMS 4999 standard, which involved annealing the samples at 538 oC +/- 14 for 4 hours +/- 0.25.

UTS (ksi) 136.5 139.8 138.2 139.7 140.9 139.9Heat Treated Wall 1 Room Heat Treated Wall 2 Room Heat Treated Wall 3 Room Heat Treated Wall 4 Room Heat Treated Wall 5 Room Heat Treated Wall 5 Room0.3 YS (ksi) 123.6 126.8 124.7 126.6 128.0 127.6Elong (%) 13 15 15 11 15 15Reduction in Area (%) 18 32 24 13 29 30Modulus (Msi) 15.5 15.6 15.6 15.9 16.1 16.0In the as-deposited condition, the samples have similar Yield Strength (YS) and ultimate tensile strength (UTS) irrespective of the processing condition involved. However, there is some scatter in the elongation to failure. The similarity of YS may arise from near similar microstructures in the as-deposited condition, whereas the scatter in elongation may result from the presence of porosity in the as-deposited samples. Upon heat treatment, there is a noticeable drop in the YS and UTS values in all cases, and yet again the properties are within scatter. There is an increase in elongation on the heat-treated samples in general.

There is an increase in elongation on the heat-treated samples in general. The reduction in strength along with the increase in elongation is indicative of microstructural coarsening that is to be expected at the annealing temperatures. To better understand the impact of varying cooling times and clamping states during deposition of these samples, the microstructures from two extreme cases (Wall 1 and Wall 4) were evaluated via scanning electron microscopy. The micrographs are presented in Figure 7.Microscopic examination shows the presence of basketweave in both the processing conditions. However, the similarity in microstructures is intriguing. These similarities in microstructures may have been caused by the start plate, or heat sink, being much larger than the part itself. Thus, Wall 1 maintained a high enough cooling rate over 10 layers to result in a basketweave microstructure.

Thus, Wall 1 maintained a high enough cooling rate over 10 layers to result in a basketweave microstructure. However, in the case of Wall 4, the cooling rate was as fast as that of Wall 1 during the deposition of the first two layers. In such a case, the deposited bead may already have cooled down to the background substrate temperature, which acts as a heat sink. On deposition of the subsequent two layers after a 30 minutes wait, if the prior layers are already at the background temperature, then the cooling rate will be similar as the build progresses. Thus, in the limited 10-layer build, the effect of the larger heat sink dominates over any potential impact the inter-layer cooling time can have.

Thus, in the limited 10-layer build, the effect of the larger heat sink dominates over any potential impact the inter-layer cooling time can have. However, for taller, it is very likely that a continuous build434may have a coarser microstructure compared to one that has substantial inter-layer cooling time because of the thermal buildup resulting from the poor thermal conductivity of titanium.Figure 7 Microstructures of Wall 1 and Wall 4 in as-deposited conditions5. Conclusions and Future Work1) Cooling time between deposited layers influences the distortion of the plate, and unclamped samples showed more distortion than the clamped samples.2) Longer inter-layer cooling time will cause larger distortion. 3) Since the built walls were not high, all the heat was transferred to the large heat sink or start plate thus, similar microstructure and mechanical properties are seen.4) Wall 1 showed minimal distortion.

Thus, for building large scale parts that are 10 layers or less, it is recommended not to use any inter-layer cooling time. This way total production time can be reduced.5) Cooling time influences distortion but not mechanical properties for large scale parts that are not high.6) There is a need to correlate amount of distortion to microstructure and mechanical properties for large scale components.7) There is also a need to focus on the deposition pulling away from the start plate due to the residual stress build up.6. AcknowledgementsThe research was sponsored by the US Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office under contract DE-AC05- 15 00OR22725 with UT-Battelle, LLC. This work is also supported in part by Cooperative Research and Development Agreement with GKN Aerospace, under contract No. NFE-15- 05725.7. References[1] B. Dutta, V. Singh, H. Natu, J. Choi, and J.

Large thermal gradients during directed energy deposition often led to residual stresses in the final deposit. Parameter optimization is predominantly used for residual stress mitigation. However, the effect of process parameters is material specific. Current research aims to study the effect of inter-layer dwell time on residual stresses in directed energy deposition of high strength steel alloy. Specimens were deposited at three levels of inter-layer dwell time. Surface as well as bulk residual stresses were measured using X-ray diffraction. Both surface as well as bulk residual stresses were found to increase with an increase in the inter-layer dwell time.Keywords: Directed Energy Deposition; Residual Stress; Inter-layer dwell time; High strength steel alloy1.

IntroductionAdditive manufacturing (AM) of metals and alloys plays a significant role in the current manufacturing scenario due to its freedom for design and lower material consumption compared to the conventional subtractive manufacturing methods. Metal AM is being adopted by various industries for aerospace, automotive, and medical applications where customized parts having complex designs are often required [1]. Extensive material development for metal AM during the last decade have resulted in more industries adopting this technology. Various steels, titanium and aluminum alloys, Nickel based super alloys as well metal matrix composites are being used for metal AM [2]. However, wider adoption of metal AM is restricted by various challenges related to the design and manufacturing of parts.

However, wider adoption of metal AM is restricted by various challenges related to the design and manufacturing of parts. Some of these challenges include porosity in the parts, lower surface finish, distortion, and poor dimensional accuracy.One of the important challenges that hinder the adoption of metal AM parts is the formation of residual stresses. Layer by layer deposition during metal AM causes cyclical heating and cooling of the metal deposit resulting in formation of residual stresses[3]. Presence of residual stresses is often undesirable as it could lead to delamination and cracking of deposited parts. Residual stresses can also cause early failure of a component, which is highly undesirable in critical applications. Different methods are adopted to mitigate residual stresses in metal AM parts. Additively manufactured metal parts are subjected to heat treatment to release the stresses, adding to the cost of manufacturing[4]. Vyatskikh et al.

Additively manufactured metal parts are subjected to heat treatment to release the stresses, adding to the cost of manufacturing[4]. Vyatskikh et al. (2023) proposed alloy design having solid state transformations with introduction of soft and hard metallic phases as a method to reduce residual stress formation [5]. Development of digital twins for in-situ monitoring and optimization of process parameters for residual stress mitigation is being investigated [6]. Parameter optimization is often used to reduce the formation of residual stresses. Knowledge of material behavior to different process parameters is required to manipulate quality of final parts [7].1824Residual stress evolution in additively manufactured parts has been studied by researchers over the years. Knowledge of influence of different process parameters on the residual stress formation is crucial to improve the properties of final deposit. Denlinger et al.

Knowledge of influence of different process parameters on the residual stress formation is crucial to improve the properties of final deposit. Denlinger et al. (2015) studied the effect of inter-layer dwell time on distortion and residual stress evolution during directed energy deposition of Ti-6Al-4V and Inconel® 625 [8]. Both alloys exhibited contrasting behavior with change in inter-layer dwell time as shown in Figure 1. Difference in the behavior was attributed to the lack of phase change in Inconel® 625 compared to BCC to HCP change in Ti-6Al-4V. Lee et al. (2023) studied the effect of inter-layer cooling on microstructure and mechanical properties of different titanium alloys. Inter-layer cooling was found to be influencing the microstructure formation and thus the properties of deposited alloys differently [9].

Inter-layer cooling was found to be influencing the microstructure formation and thus the properties of deposited alloys differently [9]. Variation of residual stresses with change in process parameters was observed to be material specific, necessitating the need to study material behaviour for new materials.Figure 1. Variation of residual stress evolution with inter-layer dwell time in DED of Titanium and Nickel alloys; with permission from publisher [8]High strength steel alloy is a recently developed low alloy steel having high strength and toughness and exhibits excellent printability [10]. As the variation of residual stresses to process parameters is material specific, current research aims to investigate the residual stress evolution in high strength steel alloy with varying inter-layer dwell time. Inter-layer dwell time can emulate the time interval between adjacent depositions at a single point of large parts.

Inter-layer dwell time can emulate the time interval between adjacent depositions at a single point of large parts. Following sections of this manuscript describes the materials and methodology of the research, experimental details, results and discussion, and conclusion.2. Materials and Methods2.1. Material CharacterizationHigh strength steel alloy powder, manufactured through gas atomization, was supplied by the Powder Alloy Corporation, USA. Elemental composition of the alloy powder as reported by the manufacturer is given in Table 1. As the quality of powder being used in the DED process affects the properties of deposited parts,1825powder characterization was conducted before sample preparation. Particle size, particle shape and presence of oxidation on the powder particles were analyzed.Table 1.

Particle size, particle shape and presence of oxidation on the powder particles were analyzed.Table 1. Chemical composition of High Strength Steel AlloyElement Iron Carbon Manganese Silicon Chromium Nickel Copper Vanadium Molybdenum Phosphorous Sulphur Aluminum Titanium Oxygen P + Sn + As + Sb HydrogenMass % Balance 0.28 0.60 1.0 2.6 1.0 <0.2 0.15 0.86 0.009 0.005 0.004 <0.006 0.03 <0.035 6 ppmShape of the High strength steel alloy powder particles was analyzed using a Scanning Electron Microscope (SEM) as shown in Figure 2a. Small, irregular-shaped satellite powder particles were observed to be attached to the large, spherical particles. Particle size distribution of the powder used for deposition was analyzed by Static Light Scattering (SLS). An average size distribution of 𝐷(cid:2869)(cid:2868) = 59 𝜇𝑚, 𝐷(cid:2873)(cid:2868) = 76.8 𝜇𝑚 and 𝐷(cid:2877)(cid:2868) = 103.2 𝜇𝑚 was obtained as shown in Figure 2b.

An average size distribution of 𝐷(cid:2869)(cid:2868) = 59 𝜇𝑚, 𝐷(cid:2873)(cid:2868) = 76.8 𝜇𝑚 and 𝐷(cid:2877)(cid:2868) = 103.2 𝜇𝑚 was obtained as shown in Figure 2b. Optical Microscopy (OM) image of the alloy powder is shown in Figure 2Figure 3. Significant oxidation of the powder particles was visible under the microscope as seen by the discoloration.Figure 2. a) SEM image of High Strength Steel Alloy powder. b) Particle size distribution of Alloy powder1826Figure 3.Optical microscope image of High Strength Steel Alloy powder2.2. Additive ManufacturingHigh Strength Steel Alloy powder was deposited onto AISI 1015 carbon steel substrates using an in- house developed Laser-DED system as shown in Figure 5. The laser-DED system employed a 1 KW Nd:YAG laser with a spot diameter of 2 mm. Argon gas was used to provide inert atmosphere in the deposition chamber.

The laser-DED system employed a 1 KW Nd:YAG laser with a spot diameter of 2 mm. Argon gas was used to provide inert atmosphere in the deposition chamber. Rectangular specimens of 14 𝑚𝑚 × 10 𝑚𝑚 cross-sectional area was deposited with a zig-zag scan pattern as shown in Figure 4. Each specimen was composed of 7 layers without any scan rotation between adjacent layers.Figure 4. Specimen deposited by Laser DEDFigure 5. Laser DED system used for deposition.1827Trial experiments were conducted to optimize the process parameters for minimum porosity and deposition geometry. Optimized parameters used for deposition are shown in Table 2. To study the effect of inter-layer dwell time on residual stress formation, three levels of inter-layer dwell time at 0 s, 10 s and 30 s were provided in between the deposition of each layer.Table 2.

Optimized Parameters used for deposition.Laser Power(W)Scan Speed(mm/min)Feed Rate(g/min)Hatch Spacing(%)Inter-layer Dwell Time (s)04505001.26010302.3. Residual Stress MeasurementHigh Strength Steel Alloy specimens prepared by directed energy deposition for this study are shown in Figure 6. X-Ray diffraction followed by uniaxial sin(cid:2870) ∅ method was used to measure the residual stresses. Initially, stresses were measured on the top surface of As-deposited specimens. To analyze the residual stress variation pattern in the bulk, specimens were mounted and polished using 1200 grit SiC paper to remove material of 500 𝜇𝑚 thickness from the top surface [5]. Even though polishing could release some of the residual stresses, objective of the study was to analyze the pattern of stress variation. Hence, same polishing conditions were maintained for all the samples. As Deposited surface and bulk surface after polishing are shown in Figure 6.Figure 6.

Hence, same polishing conditions were maintained for all the samples. As Deposited surface and bulk surface after polishing are shown in Figure 6.Figure 6. High Strength Steel Alloy specimens deposited by laser DED3. Results and DiscussionHigh strength steel alloy deposited by DED process at 3 levels of inter-layer dwell time were subjected to X-ray diffraction. Surface residual stresses of as-deposited specimens measured by uniaxial sin(cid:2870) ∅ method is shown in Figure 7. Bulk residual stresses on surface obtained by removing material of 500 𝜇𝑚 thickness from the top surface is shown in Figure 8.1828Figure 7. Variation of surface residual stresses with inter-layer dwell timeFigure 8. Variation of bulk residual stresses with inter-layer dwell timeSurface residual stresses increased with increasing levels of inter-layer dwell time.

Variation of bulk residual stresses with inter-layer dwell timeSurface residual stresses increased with increasing levels of inter-layer dwell time. Similar trend was observed for bulk residual stresses as well, reiterating the effect of inter-layer dwell time on residual stresses in high strength steel alloy deposition. As the inter-layer dwell time increases, more heat energy can dissipate from the deposited layer towards the substrate. This decreases the peak temperature in the already deposited layers. Hence, a higher thermal gradient is present when a new layer is deposited, leading to higher residual stress formation during the solidification and cooling of top layer. As the inter-layer dwell time increases, higher will be the residual stress formation, which is explained by Thermal Gradient Mechanism [11].

As the inter-layer dwell time increases, higher will be the residual stress formation, which is explained by Thermal Gradient Mechanism [11]. Similarly, compressive nature of bulk stresses can be explained by Cool-down Phase Mechanism which predicts compressive stresses in the bulk layers due to the shrinkage of top layer during cooling [12].Compressive stresses were also obtained for the top layer, which is contradicting with the common observation of tensile stresses in the top layer. This was due to a small depression in the middle of top surface where the X-ray Diffraction was performed. Residual stress formation in alloys can be affected by the variation of phase formation, owing different values of properties such as thermal expansion coefficient of formed phases1829[9].

Residual stress formation in alloys can be affected by the variation of phase formation, owing different values of properties such as thermal expansion coefficient of formed phases1829[9]. Further experiments and phase analysis are required to investigate the role of different phases being formed at different inter-layer dwell times.4.(cid:32)ConclusionThis work investigated the effect of varying inter-layer dwell time on the residual stresses of high strength steel alloy manufactured by directed energy deposition. Based on the results obtained, increasing the dwell time between layers during deposition increases the residual stresses in high strength steel alloy deposits due to higher thermal gradients. Surface residual stresses as well as bulk residual stresses varied similarly with bulk stresses having higher magnitudes due to shrinking of top layer during cooling phase.

Surface residual stresses as well as bulk residual stresses varied similarly with bulk stresses having higher magnitudes due to shrinking of top layer during cooling phase. Further investigation to identify the phase formation at different dwell times is required.5.(cid:32)AcknowledgmentThis work was partially supported by the Missouri University of Science and Technology’s Kummer Institute for Student Success, Research and Economic Development through the Kummer Innovation and Entrepreneurship Doctoral Fellowship. The support provided by DEVCOM Army Research Laboratory is appreciated.6.(cid:32)References[1]Armstrong M, Mehrabi H, Naveed N. An overview of modern metal additive manufacturing technology. J Manuf Process 2022;84:1001–29. https://doi.org/10.1016/j.jmapro.2022.10.060.[2]Bandyopadhyay A, Zhang Y, Bose S. Recent developments in metal additive manufacturing. Curr Opin Chem Eng 2020;28:96–104.

J Mater Process Technol 2015;215:123–31. https://doi.org/10.1016/j.jmatprotec.2014.07.030.[9]Lee H J, Narayana P L, Kim J H, Park C H, Hong J K, Yeom J T, et al. Effect of interlayer cooling on the microstructure and mechanical properties of titanium alloys fabricated using directed energy deposition. J1830Alloys Compd 2023;953:170202. https://doi.org/10.1016/j.jallcom.2023.170202.[10] Vaughan M W, Elverud M, Ye J, Seede R, Gibbons S, Flater P, et al. Development of a processoptimization framework for fabricating fully dense advanced high strength steels using laser directed energy deposition. Addit Manuf 2023;67. https://doi.org/10.1016/j.addma.2023.103489.[11] Mercelis P, Kruth J P. Residual stresses in selective laser sintering and selective laser melting. RapidPrototyp J 2006;12:254–65. https://doi.org/10.1108/13552540610707013.[12] Bastola N, Jahan M P, Rangasamy N, Rakurty C S.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEFFECT OF INTER-LAYER TIME INTERVAL ON THE MECHANICAL BEHAVIOR OF DIRECT LASER DEPOSITED TI-6AL-4VBrian Torries1, Shuai Shao1, Nima Shamsaei2,*, Scott M. Thompson21Center for Advanced Vehicular Systems (CAVS), Mississippi State University, MS 397622Department of Mechanical Engineering, Auburn University, Auburn, AL 36849Corresponding author: shamsaei@auburn.eduAbstractDue to its wide applicability in the biomedical and aerospace fields, where unique and/or difficult to machine geometries are required, Ti-6Al-4V continues to be a strong candidate for additive manufacturing.

In this study, the effect of inter-layer time interval on the mechanical behavior of Ti-6Al-4V fabricated via Laser Engineered Net Shaping (LENSTM) is investigated. Two sets of specimens were fabricated, each with their own inter-layer time interval, accomplished by depositing either one or two specimens per operation. Tensile tests and fully reversed, strain controlled fatigue tests were conducted on the specimens. Experimental results indicate that specimens fabricated using longer inter-layer time intervals possess a higher ultimate tensile strength, lower ductility, and finer microstructure relative to those fabricated using shorter time intervals. Additionally, specimens fabricated using longer inter-layer time intervals possessed shorter fatigue lives due to presence of more process defects, such as pores and lack of fusion, inherent to additive manufacturing.

Such effects are important to consider when producing multi- part assemblies or large parts.Keywords: Fatigue, Fractography, Tensile properties, Microstructure, Geometry effects1. IntroductionA promising avenue for manufacturing currently being researched is that of additive manufacturing (AM). Rather than traditional manufacturing, which involves subtractive machining operations, casting, or other methods for material shaping, AM has the ability to create near net shapes through repetitious deposition of material, usually in a layered fashion.

While AM can benefit a variety of applications, parts created through AM are susceptible to porosity and other defects [1], which may affect their mechanical and fatigue properties [2-4].Previous studies indicate that the localized, transient heat transfer and time variant cooling rates inherent to directed energy AM processes, such as Laser Engineered Net Shaping (LENSTM), can lead to unique microstructures vastly different from those found in wrought material [5-11].1272For Ti-6Al-4V, there is a noted tendency for the LENS material to possess the martensitic phase rather than a stable α+β phase with large lamellar, globular, or bi-modal morphologies, usually found in mill annealed wrought Ti-6Al-4V [5-11]. Some studies have reported coarse α laths, with evidence of prior β grains, near the center of Ti-6Al-4V specimens created using the LENS process [5, 12].

Some studies have reported coarse α laths, with evidence of prior β grains, near the center of Ti-6Al-4V specimens created using the LENS process [5, 12]. Even changing the inter-layer time interval of additive processes, while holding other process parameters constant, has been shown to produce different microstructural features. Yadollahi et al. [13] found that for stainless steel 316L, varying the time interval between layers for LENS fabrication resulted in different microstructures and mechanical properties. Specimens with longer inter-layer time intervals exhibited finer microstructure, lower elongation to failure, and higher strength.

Specimens with longer inter-layer time intervals exhibited finer microstructure, lower elongation to failure, and higher strength. Additionally, the specimens with longer inter-layer time intervals exhibited much higher porosity [13].Based upon the known differences in microstructure found between additive and wrought material, as well as between additive materials with different inter-layer time intervals, it is clear that an understanding of the effects on mechanical and fatigue properties of altering the inter-layer time interval while holding other process parameters constant is needed. If changing the inter- layer time interval of additive processes significantly impacts the mechanical properties of a part, then doubt is cast upon the ability of small representative specimens to accurately replicate the mechanical and microstructural of the final part with larger dimensions.

Therefore, it is vital that the effects of varying the inter-layer time interval be quantified in order to better understand AM property-performance relationships and improve the applicability of test specimen data to the design of parts intended for service.This study aims to determine the effects of varying time interval during LENS fabrication of Ti-6Al-4V, with emphasis on fatigue behavior. Therefore, two batches of Ti-6Al-4V specimens were manufactured employing different inter-layer time intervals; these specimens were then machined and annealed. Microstructures of the specimens are examined, and mechanical and fatigue properties of the specimens were evaluated. Fractography was conducted in order to determine crack initiation and propagation characteristics, as well as to examine the porosity contained within the specimens.2.

Fractography was conducted in order to determine crack initiation and propagation characteristics, as well as to examine the porosity contained within the specimens.2. Experimental SetupFor this study, thirty-six cylindrical specimens (110 mm tall, Ø = 11 mm) were fabricated using an Optomec LENSTM 750 equipped with a 1 kW Nd:YAG laser. The specimens were printed in two batches of 18 specimens; one set was printed with an inter-layer time interval of approximately 1 second, while the other set was printed with an inter-layer time interval of approximately 33.5 seconds.

This difference in time interval was accomplished by fabricating either one or two specimens at a time; therefore, the specimens with the shorter time interval are referred to as “single built specimens,” and the specimens with the longer time interval as “double built specimens.” All specimens were built upon 152.4 mm by 152.4 mm, 3.175 mm thick Ti- 6Al-4V substrates that have been clamped to the worktable at each corner. Spherical Ti-6Al-4V1273Grade 5 Plasma Rotating Electrode Process (PREP) powder with a mesh size of -100/+325, sourced form Phelly Materials Inc., was utilized during the deposition process. All specimens were fabricated in an argon atmosphere, using a laser power of 350 W, a powder feed rate of 0.16 g/sec, and a travel speed of 16.9 mm/sec. Both the layer thickness and hatch spacing were 0.51 mm. In previous research, these parameters were shown to produce nearly fully dense specimens [12].

Both the layer thickness and hatch spacing were 0.51 mm. In previous research, these parameters were shown to produce nearly fully dense specimens [12]. After the completion of each layer, the hatching orientation used for to-be-deposited tracks was rotated 90° with respect to the previous orientation. Upon the completion of the LENS process, specimens were removed from the substrate, and all but one were machined to the final geometry shown in Figure 1, in accordance with ASTM E606 [14]. Annealing Ti-6Al-4V specimens can relieve residual stresses inherent to the DLD process [15]; therefore, after machining the specimens were annealed for one hour in an argon atmosphere using a tube furnace that was preheated to 704 °C (~1300 °F). The specimens were then air cooled to room temperature via free convection. Figures 2(a) and 2(b) show examples of the specimens after deposition and machining, respectively.Figure 1.

The specimens were then air cooled to room temperature via free convection. Figures 2(a) and 2(b) show examples of the specimens after deposition and machining, respectively.Figure 1. Dimensions of cylindrical tensile fatigue specimens based on ASTM E606 (units in millimeters).(a)(b)(c)Figure 2. Examples of specimens in the (a) as-built, (b) post-machining, and (c) post-testing state.1274In order to investigate the microstructure of the single and double built LENS specimens, an unmachined specimen from each set was sacrificed. Sections were taken from the center of the specimens and subsequently cut along the radial and longitudinal planes. These sections were then hot-mounted in PolyFast and polished.

Sections were taken from the center of the specimens and subsequently cut along the radial and longitudinal planes. These sections were then hot-mounted in PolyFast and polished. An optical microscope (Zeiss Axiovert 200) equipped with a polarized light source was used to examine the microstructure.Once machined and heat treated, the gage sections of the specimens were polished to a near mirror finish using sandpaper of progressively finer grits. Tensile tests were conducted at a strain rate of 0.001 s-1 using an Instron compression/tension machine, with an Epsilon extensometer. Fully-reversed strain-controlled uniaxial tension-compression fatigue tests were performed using an MTS 810 machine and an Instron Landmark servo-hydraulic test system in conjunction with an MTS extensometer, complying with ASTM E606-92 [14]. Figure 2(c) shows an example of a tested fatigue samples.3.

Figure 2(c) shows an example of a tested fatigue samples.3. Experimental Results and Discussion3.1 MicrostructureThe microstructures of the annealed LENS Ti-6Al-4V single and double built specimens are compared in Figure 3. The microstructure from the longitudinal section of the single built specimens, shown in Figure 3(a), consists of predominately Widmanstätten microstructure combined with large α+β colonies (large white and grey features marked by red ellipses) with ultrafine α laths originating from the grain boundary α phases. This type of microstructure exists due to the decomposition of the martensitic (α’ and α”) phases as a result of the cycling heating/cooling of deposited layers during fabrication [16], while residual martensite in LENS Ti- 6Al-4V may also be possible [17].

The microstructure of the double built specimen’s longitudinal section, as shown in Figure 3(b), displays the aforementioned Widmanstätten microstructure combined with large α+β colonies and α laths originating from the grain boundary α phases, but with much finer morphologies. The size of the α laths and grain boundary α+β colonies are all smaller for the double built specimens. This change can be attributed to the different cooling rates of the specimens during fabrication; the single built specimens possessed a significantly shorter time interval between layers, leading to heat build-up and a slower cooling rate. The double built specimens, due to the large time interval between layers, cooled to a much greater extent between layers. This in turn led to a greater temperature difference between the previous and current layer during deposition, resulting in a more rapid cooling rate. This finding has also been observed and described in other works [18].

This finding has also been observed and described in other works [18]. Additionally, while the polished surfaces of the single built specimens exhibited little porosity (Fig. 2(c)), multiple lack of fusion type pores were detected on the double built microstructure specimen surfaces (Fig. 2(d)), suggesting that the double built specimens contain many more defects than the single built specimens. The higher porosity/defects for double built specimens is most likely due to a more poorly developed melt pool being utilized during deposition of the colder layer.1275Figure 3. Cross-sectional microstructure of LENS Ti-6Al-4V: (a) single built specimens, (b) double built specimens, and examples of defects found on (c) single built specimens, and (d) double built specimens.3.2 Monotonic Tensile BehaviorThe results of the tensile tests are shown in Figure 4.

While all of the specimens exhibited nearly identical behavior in the elastic region, the double built specimens possessed higher yield strengths than the single built specimens due to their finer microstructure. The single built specimens experienced at least triple the elongation to failure than the double built specimens. The single built specimens also experienced much more plastic deformation, as they underwent the majority of the elongation during plastic deformation. In contrast, the double built specimens broke at or just after the yield point, experiencing nearly no plastic deformation. This could be due to the amount and nature of the porosity found in the double built specimens. In addition to small spherical pores similar to those found in the single built specimens (Fig. 3(c)), the double built specimens also contained large lack of fusion pores, with dimensions as large as 1 mm (Fig. 3(d)).

3(c)), the double built specimens also contained large lack of fusion pores, with dimensions as large as 1 mm (Fig. 3(d)). Rather than the smooth, uniform surface of the spherical pores, as shown in Figure 5(a), the lack of fusion pores, shown in Figure 5(b), were irregularly shaped, large and flat, giving rise to a much higher stress concentration [19]. These stress concentrations were evidenced by the earlier failure and lack of ductility of the double built specimens; once the material yielded, the pore was large enough to cause rapid specimen failure.1276Figure 4. Tensile stress-strain curves for single and double built LENS Ti-6Al-4V.Figure 5. Pores from tensile fracture surfaces of (a) a single built LENS, and (b) a double built LENS Ti-6Al-4V specimen.3.3 Fatigue BehaviorFigure 6 provides the strain amplitude versus reversals to failure for the single and double built LENS Ti-6Al-4V specimens.

It can be seen that the fatigue lives of the single built specimens are longer than those of the double built specimens by approximately an order of magnitude in the long life regime. Although the fatigue lives are much closer in the short life regime, the single built specimens still exhibit longer fatigue lives. However, it is important to note that there is scatter in the fatigue lives, particularly in the high cycle fatigue (HCF) testing of the single built1277specimens, due to the existence and distribution of the process defects, such as pores, which play a more dominating role in long life regime [12] .It is interesting to note that the strain-life behavior of the double built specimens contains much less scatter than that of the single built.

This could be due to the porosity levels of the specimens; while both sets of specimens contained porosity, the double built specimens contained lack of fusion pores orders of magnitude larger than those found in the single build specimens. Additionally, there was less percent variation in pore size for the double built specimens (~50% mean deviation) than the single built specimens (~90% mean deviation). This can explain the lesser variation in plastic zone size and stress concentration of the double built specimens; corresponding to more consistent fatigue behavior.Figure 6. Comparison strain-life fatigue behavior for single and double built LENS Ti-6Al-4V.Fractography analysis of the LENS specimens demonstrated that both sets displayed fracture surface features significantly different that those found on wrought material.

Rather than regions of clearly defined crack initiation, crack propagation, and final fracture, the single built LENS fracture surfaces, as shown in Figure 7(a), were rough and uneven, with small amounts of crack propagation and with pores, usually close to the outer surface of the specimen, serving as crack initiation sites. In contrast, the double built fracture surfaces, shown in Figure 7(b), were smoother, with cracks initiating at large lack of fusion pores, and very little evidence of crack propagation, most likely due to the lack of ductility in these specimens. This supports the argument that the double built specimens possessed much less ductility due to having lack of fusion pores and finer microstructure. Fractography also revealed that the double built specimens possessed large lack of fusion pores and a greater number of spherical pores, with a larger average size, than the single built tests.

Fractography also revealed that the double built specimens possessed large lack of fusion pores and a greater number of spherical pores, with a larger average size, than the single built tests. This supports the observation made during the microstructure investigation that the double built specimens were more porous than the single built specimens. The orientation of these larger pores in relation to the load axis created a larger projected area on the plane normal to the loading axis than those formed by the smaller, spherical pores, leading to greatly increased1278stress concentrations. Research has shown that these larger projected areas can have more detrimental effects on specimen fatigue life [20]. This larger effective area, along with the finer microstructure, could be the cause of the shorter fatigue lives in the double built specimens.

This larger effective area, along with the finer microstructure, could be the cause of the shorter fatigue lives in the double built specimens. Additionally, there was evidence of multiple crack initiation points on some of the single built fracture surfaces in short life regime, as shown in Figure 7(a). This may be explained by more ductility of single built specimens, and therefore, more plastic deformation in short life regime. This will cause the specimens to spend more time in the crack propagation stage of fatigue; thus, providing more opportunities for other cracks to initiate and grow [21, 22]. It is interesting to note that the large lack of fusion pores in the double built specimens contained multiple partially melted powder, as seen in Figure 7(b).Figure 7. Representative fatigue fracture surfaces for (a) single built and (b) double built LENS Ti-6Al-4V specimens.12794.

Representative fatigue fracture surfaces for (a) single built and (b) double built LENS Ti-6Al-4V specimens.12794. ConclusionsThe goal of this study was to quantify the effects of inter-layer time interval during additive manufacturing on the mechanical behavior of Ti-6Al-4V. It was found that the variation in the inter-layer time interval leads to different microstructures, types (i.e. spherical or lack of fusion) and levels of porosity in the build, which in turn results in different mechanical behavior. In particular, the following conclusions can be made based on this study:1. Longer inter-layer time intervals create lower bulk part temperatures and higher temperature gradients between layers, leading to higher instantaneous and more sustained/continuous cooling rates. This led to the specimens with longer inter-layer time intervals (double built specimens) possessing higher porosity/defects and finer microstructure in comparison to the single built specimens.2.

This led to the specimens with longer inter-layer time intervals (double built specimens) possessing higher porosity/defects and finer microstructure in comparison to the single built specimens.2. Of tested specimens, those with larger inter-layer time intervals possessed higher yield strength and drastically reduced ductility, with failure occurring at or shortly after yielding, due to their finer microstructure and existence of many lack of fusion pores.3. Specimens produced with an increased inter-layer time interval (double built specimens) exhibited much shorter fatigue lives than those produced with a smaller inter-layer time interval (single built specimens), due to the different microstructure associated with large lack of fusion pores resulting from the different thermal histories.4. The single built specimens exhibited more scatter in fatigue life than the double built specimens.

The single built specimens exhibited more scatter in fatigue life than the double built specimens. This appears to be due to the greater relative variation in pore size found in the single built specimens as compared to the double built specimens.5. Any changes to the inter-layer time interval or geometry employed for a build, without making a corresponding adjustment to the process parameters, can lead to significant variation in specimen porosity, microstructure, and mechanical behavior.This investigation further demonstrates that one can achieve significantly different microstructural, tensile and fatigue properties in LENS Ti-6Al-4V parts by altering the inter-layer time interval; while maintaining a constant set of process parameters. In other words, changes in manufacturing strategy, i.e.

In other words, changes in manufacturing strategy, i.e. geometry changes or increasing the amount of parts fabricated per build, can result in different, and possibly undesired, mechanical properties, casting doubt upon the representability of specimen properties in relation to the component-level performance (i.e. property-performance relationships). Proper adjustments (i.e. scaling) of processing parameters, such as laser power, scan speed, power feed rate are necessary in order to yield consistent microstructures and mechanical properties for AM parts with different manufacturing strategies.AcknowledgementResearch was partially sponsored by the Army Research Laboratory and was accomplished under Cooperative Agreement Number W911NF-15-2-0025. The views and conclusions contained1280in this document are those of the authors and should not be interpreted as representing the official policies, either expressed or implied, of the Army Research Laboratory or the U.S. Government.

Government. The U.S. Government is authorized to reproduce and distribute reprints for Government purposes notwithstanding any copyright notation herein. This manuscript was prepared while Nima Shamsaei and Scott Thompson were Assistant Professors at Mississippi State University.References[1] Sterling, A. J., Torries, B., Seely, D. W., Lugo, M., Shamsaei, N., Thompson, S. M., “Fatigue behavior of Ti–6Al–4V alloy additively manufactured by laser engineered net shaping,” in: Proceedings of AIAA SciTech 2015 Kissimmee, USA, 2015.[2] Pilchak, A. L., Bhattacharjee, A., Williams, R. E. A., Williams, J. C., “The effect of microstructure on fatigue crack initiation in Ti–6Al–4V,” in: Proceedings of International Conference on Fracture (ICF12), Ottawa, Canada, 2009.[3] Pilchak, A. L., Nakase, K. I. I., Shirai, Y., Rosenberger, A. H., Williams, J.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEFFECT OF LASER ADDITIVE MANUFACTURING ON MICROSTRUCTURE EVOLUTION OF INOCULATED Zr47.5Cu45.5Al5Co2 BULK METALLIC GLASS MATRIX COMPOSITESMuhammad Musaddique Ali RAFIQUE1, Milan BRANDT2, Mark EASTON3, Dong QIU4, Ivan COLE5, Sabu JOHN61,3,4School of Engineering, RMIT University, City Campus, Carlton, VIC 3053, Australia2Additive Manufacturing Precinct, RMIT University, City Campus, Carlton, VIC 3053, Australia 5ECP Director, Advanced Manufacturing and Fabrication, Enabling Capability Platform, RMITUniversity, City Campus, Melbourne, VIC 3000, Australia 6School of Engineering Cluster, Manufacturing, Materials and Mechatronics, RMIT University,Bundoora Campus, Bundoora VIC 3083, AustraliaAbstractBulk metallic glass matrix composites are advocated to be material of future owing to their superior strength, hardness and elastic strain limit.

However, they possess poor toughness which makes them unusable in any structural engineering application. Inoculation has been used as an effective means to overcome this problem. Zr47.5Cu45.5Al5Co2 bulk metallic glass matrix composites (BMGMC) inoculated with ZrC have shown considerable refinement in microstructure owing to heterogeneous nucleation. Efforts have also been made to exploit modern laser-based metal additive manufacturing to fabricate BMGMC parts in one step. However, the effect of laser treatment on inoculated material is unknown. In this study, an effort has been made to apply laser based additive manufacturing on untreated and inoculated BMGMC samples. It is observed that laser treatment not only refined the microstructure but resulted in change of size, morphology and dispersion of CuZr B2 phase in base metal, heat affected zone and fusion zone. This effect is documented with back scattered electron imaging.

This effect is documented with back scattered electron imaging. This provides a basis for further research to quantify this phenomenon and full-scale part development.Keywords: Additive, bulk metallic glass, composites, inoculation, phaseCorresponding author: Muhammad Musaddique Ali RAFIQUE (ali.rafique@hotmail.com, s3469212@student.rmit.edu.au)(cid:28)(cid:26)(cid:20)IntroductionOver time, there has always been a quest for hard, strong and tough material [1, 2] which can effectively withstand the loads encountered in practical extreme engineering problems. Various efforts have been made to achieve this [3, 4] and many new materials were developed. Although they were discovered in 1950s [5], very recently, bulk metallic glasses [6-10] have emerged as potential candidates to solve this problem [11].

Although they were discovered in 1950s [5], very recently, bulk metallic glasses [6-10] have emerged as potential candidates to solve this problem [11]. They possess high hardness [12, 13] and strength due to parent glassy structure but have poor ductility, hence poor toughness and exhibit catastrophic failure under the action of external load [14-16]. This problem limits their application. Efforts have been made to overcome this problem by introducing crystallinity in the glassy matrix. This may be achieved by various means including; external [17-19] introduction of reinforcing particles (ex situ) or internal (in-situ) [20-26] precipitating of same during processing (solidification [27-29], devitrification [30-33], powder metallurgy [34], foaming [35] or solid-state processing). However, this quest is still far from exhausted.

However, this quest is still far from exhausted. Narrowing down the approach, In-situ introduction of precipitates during solidification have proved out to be the best mechanism to tackle this problem and newer methods have been introduced to achieve this mostly by control of melt composition [36], chemistry, and adjustment of processing conditions [37, 38]. Successful introduction of precipitates to form a composite structure was reported for the first time by Prof. Johnson’s group at Caltech in 2000 [39]. Since then many groups in the world have produced a multitude of composites using similar philosophy [21, 28, 40-56]. As for the technique, very recently, additive manufacturing has emerged as new innovative and competitive route for a quick, efficient and one step solution to produce near net shape complex parts [57-61].

Components from nearly all types of metals and alloys can be produced by this technique (laser [58-60, 62, 63] or electron beam [64, 65] based) as the temperature obtained is a function of laser power and can be flexibly controlled over a wide range. Seeing its potential, need and interest has also sparked in the use of this technique for microstructure control [54, 66-68], manipulation [69, 70] and manufacturing of parts by bulk metallic glasses and their composites [55, 67, 71-73]. Both simulation [55, 63, 65, 74-80] and experimental [64, 67, 68, 81, 82] methodologies have been rigorously applied to study, determine and ascertain the effect of various material and process parameters on final part quality and its properties.

Different AM variants such as laser surface remelting [38, 54, 66-69, 83-89], 3D printing [90-94] LENS [95], welding [96-98], pulsing [99, 100], shock peening [101], and laser solid forming [66, 69, 99-103] have been successfully applied to fabricate parts of various types, size, nature and geometries. Recently, an established grain refinement technique – inoculation [3, 104, 105] has been applied to Zr based bulk metallic glass matrix composites and found to successfully improve properties of as cast composites [28, 106]. However, the effect of the incidence of a high energy laser with varying scan speeds on these inoculated composites is not known – an important parameter for full scale part development.

However, the effect of the incidence of a high energy laser with varying scan speeds on these inoculated composites is not known – an important parameter for full scale part development. Gaps exist in the literature and reported research in how these versatile materials whose grain structure has been improved by inoculation will behave under intense source of heat energy (laser)? What will be the effect on type, size, distribution and morphology of phases as well as size, dimensions and geometries of associated melt pools, heat affected zones and unaffected base material? The present study is aimed to bridge this gap. A systematic study is carried out in which bulk metallic glass matrix composites with different percentages of ZrC subjected to a laser source of constant size (spot size).

A systematic study is carried out in which bulk metallic glass matrix composites with different percentages of ZrC subjected to a laser source of constant size (spot size). The effect of laser power and scan speed are evaluated.972Experimental procedureAs cast bulk metallic glass matrix composites were produced using vacuum arc melting and suction casting button furnace. They were produced by mixing appropriate amounts of pure elements Zr, Cu, Al and Co with different percentages of ZrC inoculant such that a small button is produced which was placed in the hearth of furnace and melted with the help of an arc. The melted volume was placed on a small orifice and sucked through a chute into a copper mold beneath using negative pressure generated by a vacuum pump. The sucked volume adopted the shape of copper mold and generated a gradient of microstructure across its length due to the difference in cooling rate.

The sucked volume adopted the shape of copper mold and generated a gradient of microstructure across its length due to the difference in cooling rate. Resulting cast parts were cut into different small pieces and subjected to laser additive manufacturing on Trumpf® laser additive manufacturing machine with different parameters. Laser additive manufacturing: Samples were placed in the enclosed chamber and machine was operated in a condition of over pressure of inert gas. Three tracks were produced on each of three samples. The first one was made with 500 W power, 0.61 mm spot size and 600 mm / minutes scan speed, second one was made with 300 W power, 0.61 mm spot size and 600 mm / minutes scan speed while third one was made with 300 W power, 0.61 mm spot size and 1500 mm / minutes scan speed. In all three cases, the spot size was kept constant while the power and scan speed were varied to check their effect on final microstructure.

In all three cases, the spot size was kept constant while the power and scan speed were varied to check their effect on final microstructure. An increase in laser power was assumed to exert more energy input, cause more turbulence in the melt pool, increase spatter [3], denudation zones [55, 76], evaporation and distort the surface structure with irregularities (shrinkage, segregation, lapping, cracks, porosity, pin and gas holes [3, 104]) produced during solidification [4, 105, 107, 108]. High laser power is also proposed to induce solid state transformations [109] in solidified melts. These defects are even more pronounced at varying scan speeds and produce range of melt pool geometries [64], shapes and morphologies with a multitude of solidification microstructures. After laser surface treatment, samples were prepared for metallographic examination.

After laser surface treatment, samples were prepared for metallographic examination. Sample preparation: For metallographic sample preparation, the laser treated half wedge and strip were further cut at 2/3rd of distance from tip so that the area containing composite structure and effect of laser on it were exposed. Test coupons of 35 mm length, 18 mm width and 2 mm thickness were mounted with the exposed surface facing downward in 25 grams Stycast epoxy resin with 25 grams Buehler conductive filler. Mounts were cured at room temperature for 8 hours. Once cured, they were placed in Thermolyne Type 48000 furnace at 50oC for 2 hours to harden the final epoxy. Cured mounts were subsequently subjected to grinding and polishing. Grinding was performed using silicon carbide papers to 240 grit and the samples were polished using 9 μm, 3 μm and 1 μm Diamond Duo polishing solution. Final polishing was done by employing 0.04 μm colloidal silica solution for 5 minutes.

Final polishing was done by employing 0.04 μm colloidal silica solution for 5 minutes. Scanning electron microscopy: Back scattered electron microscopy of laser surface treated inoculated samples was performed on a JOEL 7001F FE-SEM equipped with thermally assisted Schottky type field emission gun at The University of New South Wales, Sydney, Australia. Microscope was operated at cathode voltage of 15 KV. Sample was placed in chamber at a vacuum of 10-6 mbar created by the diffusion pump connected to it. Secondary electron imaging was performed by a collection of electrons at wall mounted secondary electron (Everhart Thornley type) detector (not shown). Back scattered electron imaging was carried out by back scattered electron detector. These electrons carry information about compositional contrast and clearly identify areas where microstructure is 100%973glass, affected by heat (heat affected zone) and remain unaffected and unaltered (base metal).

Working distance was maintained at 10.7 ± 0.1 mm.Results and discussionElectron microscopy in back scattered imaging mode was performed in all three zones (base metal, heat affected zone and fusion zone) for each and individual track of laser on all three (untreated, inoculated with 0.25% and 0.5% ZrC) samples.Zr47.5Cu45.5Al5Co2 (inoculant = 0%) Track 1This track was laid with following parameters; Power 500 W, Spot size 0.61 mm, Scan speed 600 mm / min. Three distinct regions may be identified in micrographs of cross section of sample taken at low magnification (Figure 1 (a)). These were labelled as A, B and C which may be attributed to fusion / melt zone, heat affected zone (HAZ) and untreated / base metal zone respectively. For this sample with given parameters, a very large, wide and deep fusion zone was formed. This is due to very high power of laser which caused localised concentration of energy resulting in incipient fusion.

This is due to very high power of laser which caused localised concentration of energy resulting in incipient fusion. At the same power of laser complete melting and rapid solidification of CuZr B2 [20] occurred. This can be observed by the appearance of 100% glassy structure in region A (Figure 1(b)). Another very important characteristic of this zone was formation of spherulites. These were consistent with observations made in a previous study [54] and were in contradiction to nanocrystalline phases observed upon isothermal annealing [110, 111]. Due to the rapid heat extraction and dissipation, these non-equilibrium products were formed. The width of melt pool or fusion zone was approximately 600 μm at the surface while its depth was 300 μm. There was shrinkage of approximately 30 – 40 μm at the top in the middle.

The width of melt pool or fusion zone was approximately 600 μm at the surface while its depth was 300 μm. There was shrinkage of approximately 30 – 40 μm at the top in the middle. Corners or edges of the melt pool were also blurred or torn apart and showed signs of deterioration due to intense heat generated by highly localised source of heat (laser). As the point of observation moved away from region A, a region which was depicted by mixed or hazy areas was observed. This is zone where incomplete transformation occurred, and the structure tends to freeze or coexist as crystalline metastable phase and glassy matrix. This happened as insufficient time was given to allow for complete dissolution of solid to liquid and then from liquid back to solid. This was known as the heat affected zone (HAZ). As the name suggest, it was region which was affected by heat, but remained incompletely transformed. Some phases or features were transformed, while others remained untransformed.

As the name suggest, it was region which was affected by heat, but remained incompletely transformed. Some phases or features were transformed, while others remained untransformed. It appeared as a distinct region in this sample. The width of this region ranged from 120 – 140 μm. Finally, as the point of observation moved further deep into sample, the base metal or untreated sample was observed (Figure 1 (d)). This was characterised by observation of well-known and documented spheroidal CuZr B2 phase [20]. There were few small brittle Al2Zr fcc [112] dendrites emerging from background of glassy liquid. Phases formed were indicated by light colour while glassy matrix appeared as dark background.

There were few small brittle Al2Zr fcc [112] dendrites emerging from background of glassy liquid. Phases formed were indicated by light colour while glassy matrix appeared as dark background. Evidence of small porosity was also observed which appeared as areas of black colour indicating empty space.974Figure 1 (a) – (d): Zr47.5Cu45.5Al5Co2 (0% inoculant) Track 1, (a) low magnification (x50) image of cross section of laser track, (b) Back scattered electron image of fusion zone / melt pool indicating formation of glass and spherulites avoiding nanocrystallites, (c) back scattered electron image of heat affected zone and (d) back scattered electron image of base metal indicating presence of different phases (CuZr B2 and small Al2Zr fcc phase [112])Zr47.5Cu45.5Al5Co2 (inoculant = 0%) Track 2This track was produced with following parameters; Power 300 W, Spot size 0.61 mm, Scan speed 600 mm / min.

Once again, three distinct regions may be identified in micrographs of cross section of sample taken at low magnification (x50) (Figure 2 (a)). Once again, these were labelled as A, B and C which were representation of fusion / melt zone, heat affected zone (HAZ) and untreated / base metal zone respectively. For this sample, with given parameters, a large, wide and deep fusion / melt zone was observed. This again was due to the high power of the laser which resulted in concentration of energy causing incipient fusion and formation of melt pool. As the spot size and scan speeds were kept constant while laser input power was reduced, it resulted in less deep melt pool. Also, there was less evidence of turbulence caused by reduced laser power. There was975some blurring or inconsistency at the top of melt pool but still, it was not as evident as was the case when laser power was 500 W. Some cracking was also observed at the edges of heat affected zone (HAZ).

Some cracking was also observed at the edges of heat affected zone (HAZ). This was due to the difference in temperatures between the melt pool and base metal which caused thermal stresses. Another reason for this may be attributed to emergence of different coefficients of thermal expansion in metal solidified at melt pool (glass) and the composite structure of base metal. Another phenomenon of interest observed in this track was the emergence of small precipitates in the melt pool region (Figure 2(b)). These appeared in the form of small equiaxed grains. These may be an indication of incomplete fusion due to the low energy input. Another reason for this may be attributed to complete melting due to intense localised heating, freezing and then precipitation of large number of small equiaxed grains from the glass (recrystallisation / devitrification) because of heating below glass transition temperature (Tg). There was no evidence of change in morphology.

There was no evidence of change in morphology. Another possibility which may be considered was complete fusion and then CuZr B2 spheroidal phase precipitation from the rapidly solidifying melt. Their size was small as rate of heat extraction from the very small area was very fast which resulted in nucleation of large numbers of small grains because of suppressed kinetics. As the point of observation moved away from melt pool / fusion zone, a profound heat affected zone (HAZ) was observed in the form of boundary / interface between melt pool and base metal. This region976(e)Figure 2 (a) – (e): Zr47.5Cu45.5Al5Co2 (0% inoculant) Track 2: (a) Low magnification (x50) back scattered electron image of cross section of laser track.

This region976(e)Figure 2 (a) – (e): Zr47.5Cu45.5Al5Co2 (0% inoculant) Track 2: (a) Low magnification (x50) back scattered electron image of cross section of laser track. (b) Melt pool and lower region of heat affected zone (x500), (c) Upper region of heat affected zone and its interface with base metal (d) Back scattered electron image of base metal indicating presence of CuZr B2 and small Al2Zr fcc phase [112], (e) Base metal at higher magnification (x2500) indicating presence of glassy matrix (dark grey areas) and ductile precipitates (light areas).was characterised by emergence of crystalline phases and a glassy matrix. However, as the laser power was low, its effect was not as profound as was when power was 500 W. Small cracks were also observed at the edges of the HAZ with base meal. The size of this region was around 120 – 140 μm.

Small cracks were also observed at the edges of the HAZ with base meal. The size of this region was around 120 – 140 μm. Once again, the reason for the cracking may be attributed to difference of temperature or coefficient of thermal expansion between different phases of material.

Regular spheroidal shape CuZr B2 [20] precipitates were observed along with tree / flower like small brittle Al2Zr fcc phase [112] as the point of observation moved further deep into sample towards base metal zone.977BZr47.5Cu45.5Al5Co2 (inoculant = 0%) Track 3AFigure 3 (a) – (d): Zr47.5Cu45.5Al5Co2 (0% inoculant) Track 3: Power 300 W, Spot size 0.61 mm, Scan speed 1500 mm / min, (a) Low magnification (x50) back scattered electron micrograph of cross section of laser track (b) Melt pool / fusion zone and lower region of heat affected zone (HAZ) at higher magnification indicating complete melting, solidification and 100% glassy structure, (c) Upper region of heat affected zone and its interface with base metal (d) Back scattered electron image of base metal indicating presence of CuZr B2 and small Al2Zr fcc phase [112].

Different morphologies of phases can be observed as light areas originating from back ground of glassy matrix (dark areas).CThis track was laid out with following parameters; Power 300 W, Spot size 0.61 mm, Scan speed 1500 mm / min. As the power and spot size were kept constant while scan speed was increased, it resulted in shallower and narrower melt pool or fusion zone as laser tends to move quickly in a given time over an area causing less abrupt action. Consequently, less volume was melted, and978only surface effects were observed. There was almost no turbulence and no pronounced effect of shrinkage was observed. Absence of turbulence also promoted homogenisation and fully glassy structure was formed without shrinkage, segregation and appearance of defects caused by movement of melt in confined volume, its interaction with environment (gases and air) and its entrapment. Size of melt pool was approximately 600 μm wide and 350 μm deep.

Size of melt pool was approximately 600 μm wide and 350 μm deep. There was small localised depression in the middle of pool which indicated localised shrinkage. As very less volume was melted and the interaction of energy with matter was extremely brief, size of heat affected zone was subsequently decreased and it spaned to a maximum width of 100 μm. Considerable decrease in size of heat affected zone (HAZ) indicated effectiveness and suitability of process to develop layer by layer (LBL) patterns as the need of removal of excess unwanted hard material was minimised. Grains in base metal region were mostly uniformly spaced. CuZr B2 tends to retain its near spheroidal morphology but was observed in sparse locations.

Grains in base metal region were mostly uniformly spaced. CuZr B2 tends to retain its near spheroidal morphology but was observed in sparse locations. Most of microstructure was filled with three dimensional small brittle Al2Zr fcc phase [112] appearing in the form of randomly but uniformly distributed dendrites whose size ranged from 10 μm to 25 μm.Zr47.5Cu45.5Al5Co2 (inoculant = 0.25%) Track 1Track produced by incidence of laser on inoculated material surface, subsequent analysis of zones developed, and microstructure evolved in these were explained in figure 4 (a) - (d). These represent areas where laser energy produced different features on surface of inoculated material. These materials were characterised by presence of certain amount of ZrC inoculant. This imparted grain refinement to material and decreased the size of features (primarily CuZr B2) which can be easily observed in back scatter electron images of base metal (Figure 4 (d)).

This imparted grain refinement to material and decreased the size of features (primarily CuZr B2) which can be easily observed in back scatter electron images of base metal (Figure 4 (d)). However, effect of laser generated different regions which will be explained in detail below. Laser parameters on this sample were; Power 500 W, Spot size 0.61 mm, Scan speed 600 mm / min. Figure 4 (a) showed general form and shape of melt pool. It indicated that melt pool was deep and narrow. This was due to intense heat generated by high powered laser localised in a confined area. This was backed up with a small spot size and slow scan speed which helped in further confinement of heat in small area and increased the depth to which it could penetrate. Overall width of melt pool and its depth was high as compared to width and depth when material of same composition without inoculants was subjected to laser surface treatment.

Overall width of melt pool and its depth was high as compared to width and depth when material of same composition without inoculants was subjected to laser surface treatment. Width of melt pool ranged to maximum of 900 μm while its depth extended roughly to 850 μm. Structure of melt pool showed complete fusion, dissolution of all precipitates, homogenisation / mixing, rapid cooling and formation of 100% glassy structure. There was marked indication of cracking due to intense heat generated in a confined area and sharp temperature gradient which existed between fusion zone and neighbouring cold base metal. Cracking may have also been caused due to frictional forces generated during grinding as the material was hard at fusion zone as compared to surroundings and this must have exhibited resistance to absorb energy and cracked. This crack extended all throughout the volume of material starting from surface towards base metal.

This crack extended all throughout the volume of material starting from surface towards base metal. Soon after fusion zone, an area depicted by intense effect of heat was witnessed. This was identified as heat affected zone (HAZ). This was area which was marked by incomplete / partial transformation. Size of this heat affected zone was very pronounced and profound in this sample. It ranged to maximum of 200 μm wide to 200 μm deep / long. An important feature was observed in this sample. This was presence of small crystals in partial or incompletely transformed glassy matrix. Appearance of these crystals may also be attributed to979inoculation effect itself. As there were stable inoculants present in melt which provided potent sites for heterogeneous nucleation, this sample showed appearance of small crystals even in heat affected zone. Presence of small hairline cracks indicated development of air pockets in solidifying metal. These may have also been caused by shrinkage.

Presence of small hairline cracks indicated development of air pockets in solidifying metal. These may have also been caused by shrinkage. Finally, as the point of observation moved away from HAZ towards base metal, microstructure consisting of small precipitates (mainly CuZr B2) emerging out of glassy matrix was observed. This microstructure was characterised by finely dispersed crystals which were present all throughout the matrix. Size of these precipitates / crystals was on average around 2 – 6 μm while their morphology was spheroidal. They were dispersed all throughout the volume of material. They became the source of increase in ductility and toughness of material. There were other small precipitates present in matrix as well bearing dendritic structure which were small brittle Al2Zr fcc phase [112]. They were consistent with similar structure observed in first glass matrix composite produced by Prof.

They were consistent with similar structure observed in first glass matrix composite produced by Prof. Johnson’s group [39].Zr47.5Cu45.5Al5Co2 (inoculant = 0.25%) Track 2As the laser power was reduced while keeping spot size and scan speed constant, structures shown in figure 5 (a) – (e) appeared. These clearly showed the effect of decrease in laser power which generated less heat and thus caused melting of less amount of material. This in turn generated small melt pool whose size was maximum 800 μm wide and 450 μm deep. Decrease in penetration depth was an indication that low power of laser did not generated enough heat at a point to cause deep impression. However, 100% glassy structure was formed and observed in melt pool which was an evidence that amount of heat was enough to cause complete fusion followed by rapid solidification. An important feature shown and observed in this sample was presence of small region in which no melting has occurred.

An important feature shown and observed in this sample was presence of small region in which no melting has occurred. This was characterised by a dark colour spot (Figure 5 (b)). This may have been produced because of unmelted ZrC inoculant particle at the centre of growing grain. The grain itself had not grown and may have been frozen due to extremely high cooling rate in a confined volume at the corner of small melt pool. There are few dark spots at top right corner of micrograph which indicated presence of porosity or gas holes.

There are few dark spots at top right corner of micrograph which indicated presence of porosity or gas holes. Once again, these may have been caused by poor melt treatment, improper casting practice or improper sample preparation.980Figure 4 (a) – (d): Zr47.5Cu45.5Al5Co2 (0.25% inoculant) Track 1: Power 500 W, Spot size 0.61 mm, Scan speed 600 mm / min981(e)Figure 5 (a) – (e): Zr47.5Cu45.5Al5Co2 (0.25% inoculant) Track 2: Power 300 W, Spot size 0.61 mm, Scan speed 600 mm / minAs the point of observation moved away from melt pol, once again, an area bearing the effect of intense heat was observed, known as heat affected zone (HAZ) (Figure 5 (c)). Once again, due to prior inoculation it showed evidence of heterogeneous nucleation in which ductile crystalline particles tend to appear randomly in matrix despite of immense cooling rate observed there. Size of this zone was also very large spanning from approximately 200 μm wide to 250 μm deep.

Size of this zone was also very large spanning from approximately 200 μm wide to 250 μm deep. It was marked by a lower and an upper line. Mostly, equiaxed grains were observed in this region with a mix of glassy matrix. However, few small columnar grains were also observed at the interface of HAZ and base metal. They were formed because of difference of temperature between these two regions and heat transfer occurring as a result (Figure 5 (d)). Finally, as the point of observation moved to base metal, a microstructure independent of any impulse was witnessed. This was glassy matrix in which number of small CuZr B2 phases were uniformly distributed. Size of these crystalline phases was clearly decreased while their amount has increased. On average, their size was found to be around 1.8 – 2.5 μm.

Size of these crystalline phases was clearly decreased while their amount has increased. On average, their size was found to be around 1.8 – 2.5 μm. This was evidence of increase in toughness of these alloys which was observed in another study based on microhardness testing described by author elsewhere [113].Zr47.5Cu45.5Al5Co2 (inoculant = 0.25%) Track 3This track was laid out with following parameters; Power:300 W, Spot size: 0.61 μm, Scan speed: 1500 mm / minutes. Effect of this increased roaster / scan speed at power of 300W can easily be seen in development of small, shallow and narrow melt pool. Like previous sample, area exhibited by base metal consisted of small equiaxed precipitates originating from background of glassy matrix (Figure 6 (d)). These were mostly equiaxed in nature dispersed evenly all throughout the matrix. Melt / fusion zone on the other hand was manifested by complete fusion and formation of glass (Figure 6 (b)).

These were mostly equiaxed in nature dispersed evenly all throughout the matrix. Melt / fusion zone on the other hand was manifested by complete fusion and formation of glass (Figure 6 (b)). There was no evidence of nucleation and growth or precipitation. Area known as heat affected zone (HAZ) which occurred in the middle of glassy matrix and base metal was shown in Figure 6 (c). Unlike previous heat affected zone (HAZ), the size of this area was reduced to width of almost 200 μm and depth of 180 μm. Primary reason of this was increased scan speed which decreased the time laser spent on metal surface thus created very short incipient melt pool982and associated heat affected zone. Morphology of crystallites in this region was again not quite well developed. However, presence of small Al2Zr fcc phase [112] can easily be witnessed by appearance of dark areas (near spheroids or sharp needle like precipitates) emerging from the background of glassy matrix.

However, presence of small Al2Zr fcc phase [112] can easily be witnessed by appearance of dark areas (near spheroids or sharp needle like precipitates) emerging from the background of glassy matrix. This area was mostly marked by presence of metastable phase. There was almost no evidence of cracking, porosity, pin or gas holes and microstructure mostly showed uniformity.Figure 6 (a) – (d): Zr47.5Cu45.5Al5Co2 (0.25% inoculant) Track 3: Power 300 W, Spot size 0.61 mm, Scan speed 1500 mm / minZr47.5Cu45.5Al5Co2 (inoculant = 0.5%) Track 1As the percentage of inoculant increased to 0.5%, a trend marked by increase in number density of phases developed was observed in as cast samples. This was optimistic indicator as a trend marked by increased toughness was expected (as reported previously [113]). However, change in type, size and morphology of same phases as affected by incident laser power was entirely different as will be explained in detail below.

However, change in type, size and morphology of same phases as affected by incident laser power was entirely different as will be explained in detail below. Figure 7 (a) showed low magnification (x50) image of melt pool when laser of 500 W stroked the sample with spot size of 0.61 mm and scan speed of 600 mm / minute. This area was marked by violent and turbulent movement of melt in confined space983when it got melted under the action of intense localised heat. There was a lot of surface roughness and sharp discontinuities on the surface. A marked depression was also observed on the surface indicating profound shrinkage. Total width / span of melt pool was calculated to be around 1000 μm at the surface while its depth was around 550 μm (highest point to bottom). The width of heat affected zone was small and may be calculated to be around 100 μm maximum. One reason which may be attributed to this was combination of speed and laser power.

The width of heat affected zone was small and may be calculated to be around 100 μm maximum. One reason which may be attributed to this was combination of speed and laser power. Also, it was anticipated that much of heat in this type of situation would be lost to atmosphere by surface exposure as melt pool was wide. When observed at higher magnification (x500), melt pool showed signs of complete melting followed by rapid solidification characterised by 100% monolithic glassy structure. No evidence of partial or complete crystallisation or recrystallisation was observed in melt pool and microstructure showed no sign of development of any feature (Figure 7 (b)). Soon after melt pool, an area marked by intense effect of heat on composite was observed known as heat affected zone (HAZ) (Figure 7(c)). This may be characterised into three regions.

Soon after melt pool, an area marked by intense effect of heat on composite was observed known as heat affected zone (HAZ) (Figure 7(c)). This may be characterised into three regions. 1) Recrystallised region (light areas) (mostly these were recrystallised equiaxed CuZr B2 grains), 2) Al2Zr fcc phase (dark grey areas) [112] and 3) Recrystallised columnar CuZr B2 grains. These are indicated by yellow arrow and found right at the interface / boundary of base metal and heat affected zone carrying equiaxed grains. Finally, when the point of observation moved further away from surface towards base metal, an area unaffected by the heat of laser was observed known as base metal. This was drastically characterised by presence of (a) small equiaxed CuZr B2 phase, glassy matrix (interdendritic dark areas) and Al2Zr fcc phase (dark grey areas (profound dendritic structure)).

This was drastically characterised by presence of (a) small equiaxed CuZr B2 phase, glassy matrix (interdendritic dark areas) and Al2Zr fcc phase (dark grey areas (profound dendritic structure)). There was also evidence of slight interdendritic shrinkage and micro porosity.984AFigure 7 (a) – (d): Zr47.5Cu45.5Al5Co2 (0.5% inoculant) Track 1: Power 500 W, Spot size 0.61 mm, Scan speed 600 mm / min.Zr47.5Cu45.5Al5Co2 (inoculant = 0.5%) Track 2Microstructure evolved under the action of decreased power of laser (300 W) with same spot and scan speed was shown in Figures 8 (a) – (e). These were characterised by presence and appearance of typical CuZr B2 phase (spheroidal or near spheroidal), Al2Zr fcc phase (sharp needle like / flower tip like dark grey areas) [112] and small porosity. Under the action of reduced laser power which generated less heat, size, shape and morphological features of melt pool were altered and were entirely different.

Under the action of reduced laser power which generated less heat, size, shape and morphological features of melt pool were altered and were entirely different. Size of melt pool was decreased, it was manifested by lesser amount of turbulence, hence less spatter, denudation zones and rough corners / edges. There was small evidence of surface depression which indicated shrinkage and small internal porosity. Overall, size / span of melt pool was reduced to 850 – 900 μm at the surface while its depth was also reduced to around 550 μm. An image of melt pool at high magnification (x500) was shown in Figure 8 (b). It clearly indicated formation of monolithic 100% glassy structure with no sign or evidence of precipitation. However, as point of observation was moved further deep into sample, an area depicted by partial melting and recrystallisation was observed. This was heat affected zone (Figure 8 (c)).

However, as point of observation was moved further deep into sample, an area depicted by partial melting and recrystallisation was observed. This was heat affected zone (Figure 8 (c)). It clearly showed sign of fusion at certain places and then recrystallisation while some areas specially elongated Al2Zr fcc phase were left as such. It appeared that CuZr B2 phase was melted completely while Al2Zr fcc phase experienced partial or incomplete melting. Size of heat affected zone was very large spanning to a depth of around 220 – 230 μm. This was shown in two parts in figures 8 (c) – (d). Base metal with unaltered structure was observed in figure 8 (e).

This was shown in two parts in figures 8 (c) – (d). Base metal with unaltered structure was observed in figure 8 (e). This was marked again with three structures shown and described previously.985B(e)Figure 8 (a) – (e): Zr47.5Cu45.5Al5Co2 (0.5% inoculant) Track 2: Power 300 W, Spot size 0.61 mm, Scan speed 600 mm / min C986Zr47.5Cu45.5Al5Co2 (inoculant = 0.5%) Track 3Finally, when the scan speed of laser was increased to 1500 mm / minute while keeping laser power at 300 W and spot size at 0.61 mm. Melt pool, heat affected zone (HAZ) and microstructures evolved in these were shown in figures 9 (a) – (d). Surface of material experienced intense effect of heat and there were clear and marked signs of surface roughness, spatter and discontinuities. However, as the laser traversed its path quickly, size of melt pool (both in terms of width and depth) was not very large. It spaned at around 70 μm on the surface while its depth was almost similar.

However, as the laser traversed its path quickly, size of melt pool (both in terms of width and depth) was not very large. It spaned at around 70 μm on the surface while its depth was almost similar. There was no indication of precipitation or recrystallization of any phase in melt pool and at higher magnification (x500) (Figure 9 (b)). Melt pool depicted placid glassy structure with no crystals. There was however, slight effect of burning at the surface which was characteristic of intense heat. Size of heat affected zone (HAZ) beneath fusion zone was not profound as well. It merely had a depth of around 120 μm. Features observed in this were fused and rapidly solidified glass and small brittle Al2Zr fcc phase. Al2Zr fcc phase showed tint of silvery or light grey areas typical of this phase as observed previously as well.

Al2Zr fcc phase showed tint of silvery or light grey areas typical of this phase as observed previously as well. Surprisingly, there was no observation of spherulites in fusion zone or HAZ of any of inoculated samples (0.25% as well as 0.5%) and it may be inferred that inoculation resisted change of morphology. Structure was found in unaltered state in base metal and found to have small to medium size precipitates around 4 – 6 μm in size. Overall effect of intense heat from localised source of energy was found to cast change in microstructure.987Figure 9 (a) – (d): Zr47.5Cu45.5Al5Co2 (0.5% inoculant) Track 3: Power 300 W, spot size 0.61 mm, Scan speed 1500 mm / min.ConclusionsIn general, following conclusions could be drawna. Inoculation treatment casted an effect of refinement on microstructure (grain size, its countand distribution).b. ZrC was found to be an effective new inoculant for bulk metallic glass matrix composites. c.

ZrC was found to be an effective new inoculant for bulk metallic glass matrix composites. c. As the percentage of inoculant increased from zero to 0.5%, type, size, shape andmorphology of phases changed, and their amount increased.d. The CuZr B2 phase becomes refined and decreases from around 4 μm to 2.8 μm with increase of percentage of inoculant from zero to 0.25% while increased slightly when percentage of inoculant increased to 0.5% indicating non-linearity.e. Another phase namely brittle Al2Zr fcc phase (small to medium columnar dendrites) wasalso observed to form out of glassy melt.f. Morphology of CuZr B2 changed from spheroids to spherulite under the effect of intenseheat observed only when laser power was 500 W and there was no inoculation.g.

Morphology of CuZr B2 changed from spheroids to spherulite under the effect of intenseheat observed only when laser power was 500 W and there was no inoculation.g. Inoculation was found to resist change of morphology of phases in melted and rapidly solidified material as no evidence of spherulites was observed in samples treated with ZrC.AcknowledgementsAuthors would like to thank Dr. Daniel East (CSIRO) for providing experimental facilities for melting and casting of samples, Dr. Dong Qiu (also author) for design of inoculants and providing their samples, Mr. Alan Jones (Additive Manufacturing Precinct, RMIT University) for performing laser surface treatment, Mr. Jeremy Gonzalez (Massachusetts Materials Research Inc., Worcester, MA) for cutting, mounting, grinding and polishing of samples and Dr. Simon Hager (The University of New South Wales, Sydney) for help in performing scanning electron microscopy in back scatter electron imaging mode.

fect of Laser Fluence in Laser-Assisted Direct Writing of Human Colon CellYafu Lin¹, Guohui Huang², Yong Huang¹*, Jeremy Tzeng², Douglas Chrisey³¹ Department of Mechanical Engineering, Clemson University, Clemson, SC 29634, USA ² Department of Biological Sciences, Clemson University, Clemson, SC 29634, USA ³ Department of Materials Science and Engineering, Rensselaer Polytechnic Institute, Troy, NY12180, USACorresponding authorABSTRACT Matrix-Assisted Pulsed-Laser Evaporation Direct-Write (MAPLE DW) has been emerging as a promising biological construct fabrication technique. The post-transfer cell viability in MAPLE DW depends on various operation conditions such as the applied laser fluence. For wide implementation of MAPLE DW-based biofabrication, the effect of laser fluence on the post-transfer cell viability should be first carefully understood.

For wide implementation of MAPLE DW-based biofabrication, the effect of laser fluence on the post-transfer cell viability should be first carefully understood. This study investigates the effect of laser fluence on the post-transfer cell viability in transferring of human colon cell HT-29. It has been observed that: 1) the HT-29 cell viability decreases from 95% to 78% as the laser fluence increases from 258 to 1,500 mJ/cm²; and 2) cell injury in this study is mainly from the process-induced mechanical stress during the cell droplet formation and cell landing processes while the effects of thermal influence and ultraviolet radiation are below the level of detection.1.

INTRODUCTION Recently, various additive direct write technologies including laser guidance direct writing [1], modified laser-induced forward transfer (LIFT) (including Matrix-Assisted Pulsed-Laser Evaporation Direct-Write (MAPLE DW) [2,3] and its modifications [4,5]), ink-jetting [6], electro-hydrodynamic jetting [7] and biopolymer deposition [8], have been investigated for living cell-based pattern and construct fabrication. Among all these technologies, MAPLE DW, a non-contact laser-based direct-write technique, has emerged as one of the most promising surface deposition and additive manufacturing techniques [9]. As shown in Fig. 1, a typical, MAPLE DW setup includes three basic components: an ultraviolet (UV) pulsed laser source; a laser transparent quartz disk coated with materials to be transferred, known as a “ribbon”; and a computer-controlled receiving substrate for the transferred material.

In next generation system the ribbon is also computer-controlled.Laser pulse Quartz support Laser pulse Quartz support Laser pulse Quartz supportLaser pulse Laser pulseRibbon Ribbon RibbonRibbon RibbonForming bubble Forming bubble Forming bubbleHuman colon Human colon Human colon cancer cells cancer cells cancer cellsForming droplet Forming droplet Forming dropletDeposited cell droplet Deposited cell droplet Deposited cell dropletReceiving substrate Receiving substrate Receiving substrateFig.1. MAPLE DW apparatus and mechanistic schematicEspecially enabling biofabrication processes like the aforementioned may introduce excessive thermal and/or mechanical stresses to biomaterials including living cells. If this process-induced stress exceeds the adaptive capacity of a cell, irreversible damage may occur. The cell injury/damage can be simply classified as the thermal and/or mechanical cell damage and the biochemical damage [10].

The cell injury/damage can be simply classified as the thermal and/or mechanical cell damage and the biochemical damage [10]. Generally, cell injury is reversible up to a certain point; however, exposure of a cell to a high magnitude and/or lasting external stress may cause irreversible cell injury even cell death. Indeed, cell death due to process-induced cell injury is common in biofabrication processes [2, 11-13], and the post-transfer cell viability is a key index to evaluate the feasibility and efficiency of a biofabrication technique.Some studies [2, 12-13] have been performed to investigate biofabrication process-induced cell injury.

For MAPLE DW, there have been some interesting contributions, which include the experimental work on the effect of Matrigel coating thickness of the receiving substrate on the post-transfer mammalian cell viability [2] and the effect of laser fluence on the post-transfer yeast cell viability [13] as well as some modeling attempts regarding the MAPLE DW process-induced mechanical stress profile during bubble expansion [14] and landing [15]. However, the effects of laser fluence on mammalian cell injury have not been documented so far as mammalian cells may be more sensitive to external stress that is largely related to the laser fluence.This study investigates the effects of laser fluence on the post-transfer HT-29 cell viability during MAPLE DW. The rest of the paper is organized as follows: First, the experimental materials and methods are illustrated in detail. The effects of laser fluence on the HT-29 cell viability are presented.

The rest of the paper is organized as follows: First, the experimental materials and methods are illustrated in detail. The effects of laser fluence on the HT-29 cell viability are presented. Finally, some conclusions on the effects of laser fluence on cell injury during MAPLE DW are drawn.2.MATERIALS AND METHODS 2.1 Cell cultureHuman colon cancer cells such as HT-29 have been widely used as a research cell source tostudy various normal and neoplastic processes [16-18].

As the HT-29 cell has retained certain characteristics of a normal tissue such as hormone receptors [19], it has been selected as a model cell to investigate the effects of laser fluence on post-transfer cell injury.The HT-29 cells (ATCC, Manassas, VA) were cultured to 90% confluency in McCoy’s 5A modified medium with L-Glutamine (ATCC, Manassas, VA), 10% (v/v) fetal bovine serum (Atlanta Biologicals, Lawrenceville, GA) and 1% (v/v) Penicillin-Streptomycin (ATCC, Manassas, VA) in a 37°C, 5% CO2, humidified incubator. Then the cultured cells were washed twice with 10 ml of phosphate buffered saline (PBS) followed by the addition of 4 ml Trypsin-EDTA (0.25% Trypsin/0.53 mM EDTA in HBSS, ATCC, Manassas, VA) and incubated for 5 minutes. The trypsinized cell suspension was transferred to a 15ml centrifuge tube and vortexed with 1ml of calf serum (Biowhitaker, Walkersville, MD).

The trypsinized cell suspension was transferred to a 15ml centrifuge tube and vortexed with 1ml of calf serum (Biowhitaker, Walkersville, MD). The suspension was centrifuged at 454 g for 5 minutes, and the resulting cell pellet was resuspended in the same medium to a final density of 1×107cells/ml for use in subsequent laser printing experiments.2.2 Laser-assisted direct writing system In order to better understand the post-transfer cell viability during MAPLE DW, direct writing of HT-29 cells have been studied using an ArF excimer laser (Coherent ExciStar, 193 nm, 12 ns FWHM duration). The MAPLE DW experimental process schematic is illustrated in Fig. 1. An Edmund optics quartz flat (Edmund optics, Barrington, NJ, 85% transmittance at 193 nm) was used as a quartz disk to make the ribbon.

1. An Edmund optics quartz flat (Edmund optics, Barrington, NJ, 85% transmittance at 193 nm) was used as a quartz disk to make the ribbon. The substrate position was computer-controlled using an Aerotech XYZ translation stage.2.3 Laser-assisted cell transfer process The supporting quartz discs and receiving substrates were washed with 70% ethanol (Fisher scientific, IL) and then rinsed with deionized water (Fisher scientific, IL). When quartz discs were dried, a 3M Scotch tape was used to make a 1.5 × 1.5 cm well with a 100 µm depth on the quartz disc. Then, 20 µl of HT-29 cell suspension was pipetted into the well, resulting in an approximately 80 ~ 100 µm thick coating. A freshly coated quartz disc was prepared for each cell transfer process.The laser spot size was focused to a 120 × 280 µm rectangle and the laser repetition rate was held at 50 Hz.

A freshly coated quartz disc was prepared for each cell transfer process.The laser spot size was focused to a 120 × 280 µm rectangle and the laser repetition rate was held at 50 Hz. Since the effect of the laser fluence on the cell viability is of interest, the laser fluence was varied from about 250 to 1,500 mJ/cm² during the experiment, and the actual laser fluence during each printing was determined based on the average of 100 laser pulses which were individually measured using a FieldMax laser power/energy meter (Coherent, Santa Clara, CA). The measured laser fluences were 258±6, 458±6, 669±10, 869±13, 1063±13, 1262±15, and 1482±15 mJ/cm².The receiving substrate was a 24-well flat bottom microliter plate and each well was filled with 1 ml McCoy’s 5A modified medium with L-Glutamine supplemented with 10% (v/v) fetal bovineserum and 1% (v/v) Penicillin-Streptomycin, resulting in a 5 mm thick cell medium coating inside each well.

The transferred cells were directly printed into the liquid media inside one of 24 wells. The cell medium in each well not only provided an impact cushion for the transferred HT-29 cells during landing, but also provided a growth medium for the transferred HT-29 cells. Three control wells each with 20 µL HT-29 cell suspension were also prepared to provide the benchmark information about the HT-29 cell viability under a nominal condition.Under each fluence level, 1000 laser pulses as a cell transfer process were applied at different locations of a newly-coated quartz disk to transfer the trypsinized cell suspension to a well. The cell transfer process was repeated three times to deposit cells into three different wells of a multiwell plate for each laser fluence. Transferred cells were tested for cell viability (as described below) immediately after printing.2.4 Cell viability testTrypan blue is a traditional and widely employed cell viability assay for HT-29 cells [20-22].

The viability of HT-29 cells was evaluated using 0.4% trypan blue stain (Biowhittaker, Walkersville, MD). Post-transfer cell suspension and trypan blue dye were mixed first at a 50%-50% volume ratio. Then the samples were viewed using an optical microscope, and the live/dead cell assay was performed using a hemocytometer. During the counting process, transparent cells were considered live while blue cells were dead since viable cells with intact cellular membrane will exclude the blue dye. Each well sample was counted twice to get an average cell viability value.3. CELL TRANSFER EXPERIMENTAL RESULT As discussed in the previous sections, the process-induced cell injury is present in different methods used in biofabrication [2,12,13].

CELL TRANSFER EXPERIMENTAL RESULT As discussed in the previous sections, the process-induced cell injury is present in different methods used in biofabrication [2,12,13]. Generally speaking, any biologically important molecule in a cell can be the target of injury producing stress, and four biochemical systems are particularly vulnerable: cell membrane and cytoskeleton system, energy metabolism, protein synthesis, and nucleic acids representing the integrity of the genetic apparatus of a cell. Since many biochemical systems of the cell are inter-dependent, injury at one site typically causes a secondary injury to other cellular processes.

Since many biochemical systems of the cell are inter-dependent, injury at one site typically causes a secondary injury to other cellular processes. Cell injury in biochemical systems is usually due to cellular stresses caused by different physical changes, and the following lists the main negative changes: 1) mechanical deformation and rupture due to crush/blast injury, fractures, lacerations, hemorrhage, osmotic shock, and noise trauma; 2) physical changes due to thermal (such as burns, heat stroke, heat exhaustion, frostbite, and hypothermia), electrical (such as electric shock), and light (such as x-ray, radioactive element, and ultraviolet radiation) damage; 3) chemical changes due to caustic agents, poisons, genotoxic and proteotoxic compounds; 4) nutritive variations due to deficiency of oxygen, vitamins and basic nutrients; and 5) biological changes due to viruses, microorganisms, protozoan and metazoan parasites) [13].

During biofabrication processes, negative changes aremainly process-induced thermomechanical stresses [13-15].There usually is a time lag between the stress and the morphologic changes during cell injury or death, and this time lag varies with the sensitivity of methods used to detect these changes [23]. Cell injury can be simply classified as the mechanical cell damage and the biochemical damage. The biochemical damage is not immediately obvious and takes a time scale of the order of a cell-generation time before changes can be detected; and the mechanical cell injury takes place instantaneously and is immediately detectable [10]. Fortunately, cells have the ability to compensate some reversible injuries through normal physiologic adaptive mechanisms. If the injury is reversible, the damaged cells can recover to their normal functions. Otherwise, the cell death happens.

If the injury is reversible, the damaged cells can recover to their normal functions. Otherwise, the cell death happens. This study has investigated the effects of laser fluence on the post-transfer cell viability and the cell injury reversibility, which may be mechanical and/or biochemical damage-based, respectively.The membrane of transferred HT-29 cells may be: lysis (broken), permeable, and intact. For lysis cells, their cell membranes are seriously damage and dissolved in the cell suspension, so they cannot be easily detected using trypan blue. Only cells with permeable and intact membranes are considered here. For intact cell membrane HT-29 cells, the blue indicator turned colorless in the presence of active enzymes, thus indicating living cells, as shown in Fig. 2. For dead/damaged cells with a cell membrane structure, their cell membranes are permeable, resulting in a blue stain inside the cells as shown in Fig.

2. For dead/damaged cells with a cell membrane structure, their cell membranes are permeable, resulting in a blue stain inside the cells as shown in Fig. 2.Post- transfer Post- transferLive cell Live cellDead cell Dead cellFig. 2. Post-transfer HT-29 cells after trypan blue treatment4. EFFECT OF LASER FLUENCE ON POST-TRANSFER CELL VIABILITYAs shown in Fig. 3, as the laser fluence increased from about 250 to 1500 mJ/cm², the HT-29 cell viability decreased from 95% to 78% (after considering the nominal apoptosis based on the control group). This relationship is mainly attributed to the process-induced mechanical stress during the cell droplet formation and cell landing processes, the process-induced thermal influence, and any residual UV radiation penetrating the cells.

To improve the cell viability during MAPLE DW, the applied laser fluence should be carefully selected.100110901008090)7080%(y t i l i b a i v6050)%( y t i l i b a i v706050l l e C4030l l e C40302020101000 Control2584586698691063Laser fluence (mJ/cm²)1262148200 Control2584586698691063Laser fluence (mJ/cm²)12621482(a)(b)Fig. 3. Post-transfer cell viability as a function of laser fluence: a) the control effect is not considered and b) the control effect is consideredThe process-induced mechanical stress during MAPLE DW may come from two different processes: the cell droplet formation (acceleration) and the cell droplet landing on the receiving substrate (deceleration). During the droplet formation, the HT-29 cell droplet is generated due to the expansion of forming the bubble, which is the result of the laser-matrix material interaction.

During the droplet formation, the HT-29 cell droplet is generated due to the expansion of forming the bubble, which is the result of the laser-matrix material interaction. The rapid expansion due to the high pressure bubble accelerates the forming cell droplet, and such an acceleration can be as high as 105~109 g and the resulting droplet velocity can be as high as 50~1000 m/s during laser-assisted cell transfer [24,25]; the higher the laser fluence, the higher pressure, resulting in a higher acceleration/velocity. The impact during the cell droplet landing process also brings a significantly higher deceleration to HT-29 cells as simulated [15,26], and the cell viability was found to be closely related to the coating thickness of receiving substrate [2]. Higher laser fluences introduce higher cell droplet accelerations during the cell droplet formation process and higher cell droplet decelerations during the cell landing process [15].

Higher laser fluences introduce higher cell droplet accelerations during the cell droplet formation process and higher cell droplet decelerations during the cell landing process [15]. As recognized, the process-induced high acceleration (or deceleration) and velocity can easily lead to severe cell injury or death as observed in the centrifugal force-induced cell damage studies [27,28]. In this study, the dependence of the cell viability on the laser fluence has proved the existence of mechanical damage as the higher laser fluence led to the higher cell acceleration (or deceleration) and velocity, resulting in the lower cell viability. It should be pointed out it is acceleration (or deceleration)-induced normal/shear stress, strain, and/or strain rate that damage cells being transferred over less than 5 microseconds [15].The process-induced thermal injury to cells or tissues might be of concern as the laser energy absorbed and vaporization/carbonization of cells.

Depending on the laser energy-induced temperature rise and its duration, thermal injury can be classified into four stages: hyperthermia, coagulation (denaturation), vaporization and carbonization [29]. In fact, thermal damage to biologicalcausesthedeactivationofenzymes,denaturationofproteins,materials is always a temperature-time governed process [30]. If a laser pulse duration is longer than 10 µs, an Arrhenius-type activation process is considered as the dominant thermal damage mechanism [31,32]. For a pulse duration less than 1 µs, the bubble formation-induced pressure is the dominant mechanism for damage [33,34]. During the 12 ns laser pulse duration in this study, the thermal penetration depth due to the Fourier heat conduction is a few micrometers as the whole cell droplet formation usually happens within a few microseconds.

Compared with the typical coating thickness (around 100 µm), the possible heat affected zone and the resulting thermal damage is considered negligible.UV radiation from laser pulses can also effectively kill living cells; fortunately, the UV light in MAPLE DW is well constrained inside a confined volume at the quartz and the coating interface as discussed below. The pulsed laser fluence threshold for photochemical DNA damage was reported as small as 4 mJ/cm² in studying the DNA damage of normal human fibroblasts (NHF) and Chinese hamster ovary cells using 193-nm radiation [35]. Using the Beer–Lambert law, the damage threshold depth is estimated just around 5 µm under a 1,000 mJ/cm² using the 4 mJ/cm²threshold laser fluence and the410 cm-1 absorption coefficient of water [36,37]. For a 100 µmthick coating in this study, it means that only 5 % of coating is affected by UV radiation (193 nm), and this should result in very limited UV radiation-related injury to cells.

For a 100 µmthick coating in this study, it means that only 5 % of coating is affected by UV radiation (193 nm), and this should result in very limited UV radiation-related injury to cells. As a matter of fact, comet assay of DNA damage showed that UV damage was not detected within the limits of the assay for MAPLE DW [2], and laser micro-dissection and laser pressure catapulting [38]. Even under high laser fluences, comet assay of DNA damage indicated that more than 97% of the specimen was not at all affected by UV radiation in laser micro-dissection and laser pressure catapulting [38]. Therefore, it is considered UV radiation-induced injury to be negligible.In this study, it is considered cell injury to be from the process-induced mechanical stress during the cell droplet formation and cell landing processes. As seen from Fig. 3, a higher the laser fluence leads to a higher pressure-induced acceleration/velocity and a concomitant lower cell viability.

As seen from Fig. 3, a higher the laser fluence leads to a higher pressure-induced acceleration/velocity and a concomitant lower cell viability. Moreover, it was also found that the post-transferred cells can proliferate through confluency as shown in Fig. 4 when incubated at 37°C, 5% CO2, humidified incubator. Future work will use more sensitive bioindicators to focus on the effects of laser fluence on post-transfer cell proliferation and cell injury reversibility of HT-29 cells during MAPLE DW.0 hour 0 hourCell Cell24 hours 24 hours 24 hoursCell Cell200 μm 200 μm200 μm 200 μm 200 μm48 hours 48 hours 48 hoursCell Cell72 hours 72 hoursCell Cell200 μm 200 μm 200 μm200 μm 200 μmFig. 4. Cell proliferation under a 258 mJ/cm2 laser fluence using a 25 laser-pulse transfer (0 to 72 hours after transfer)5. CONCLUSIONS The HT-29 cell transfer using MAPLE DW has been studied to elucidate the effects of laser fluence on the post-transfer cell viability.

CONCLUSIONS The HT-29 cell transfer using MAPLE DW has been studied to elucidate the effects of laser fluence on the post-transfer cell viability. While MAPLE DW is of interest in this study, the resulting observations are expected to be applicable to other laser-induced forward transfer studies. Some conclusions are drawn: 1) the HT-29 cell viability decreases from 95% to 78% as the laser fluence increases from 258 to 1,500 mJ/cm²; and 2) cell injury is mainly from the process-induced mechanical stress during the cell droplet formation and cell landing processes while the effects of thermal influence and ultraviolet radiation might be negligible. While this study reveals some interesting effects of laser fluence on the HT-29 cell viability, the future work will address the effects of laser fluence on post-transfer cell proliferation and cell injury reversibility of HT-29 cells using more precise bioindicators.

[27] Leverett, L.B., Hellums, J.D., Alfrey, C.P., and Lynch, E.C. Red blood cell damage by shearstress. Biophys. J. 12, 257, 1972.[28] Thoumine, O., Ott, A., and Louvard, D. Critical centrifugal forces induce adhesion ruptureor structural reorganization in cultured cells. Cell Motil. Cytoskeleton 33, 276, 1996.[29] Niemz, M.H. Laser–Tissue Interactions: Fundamentals and Applications (Springer, Berlin,1996).[30] Welch, A.J. and Van Gemert, M.J., Optical-Thermal Response of Laser-Irradiated Tissue(Plenum Press, New York, 1995).[31] Birngruber, R., Hillenkamp, F., and Gabel, V.P. Theoretical investigations of laser thermalretinal injury. Health Phys. 48, 781, 1985.[32] Sentrayan, K., Haridass, C., and Trouth, C.O. Laser induced bubble formation as a damagemechanism in the retinas of monkey and rabbit. Spectr. Lett. 34, 469, 2001.[33] Hansen, W.P., and Fine, S. Melanin granule models for pulsed laser induced retinal injury.Appl. Opt.

Spectr. Lett. 34, 469, 2001.[33] Hansen, W.P., and Fine, S. Melanin granule models for pulsed laser induced retinal injury.Appl. Opt. 7, 155, 1968.[34] Gerstman, B.S., Thompson, C.R., Jacques, S.L., and Rogers, M.E. Laser induced bubbleformation in the retina. Lasers Surg. Med. 18, 10, 1996.[35] Kochevar, I.E., Walsh, A.A., Held, K.D., Gallo, R.L., and Mirro, J. Mechanism for 193-nmlaser radiation-induced effects on mammalian cells. Radiat. Res. 122, 142, 1990.[36] Kochevar, I.E. Biological effects of excimer laser radiation. Proceedings of the IEEE 80, 833,1992.[37] Vogel, A., and Venugopalan, V. Mechanisms of pulsed laser ablation of biological tissues.Chem. Rev. 103, 577, 2003.[38] Vogel, A., Lorenz, K., Horneffer, V., Hu¨ttmann, G., von Smolinski, D., and Gebert, A.Mechanisms of laser-induced dissection and transport of histologic specimens. Biophys. J. 93, 4481, 2007..

Three different conditions including as-built and laser-polished using two different process parameters are considered. Uniaxial tension- compression fatigue tests are conducted in strain-controlled mode to examine the fatigue behavior for each condition. In addition, surface roughness measurements and fractography using optical microscopy and porosity measurements using the X-ray computed tomography are also performed for all conditions. The results indicate that laser-polishing alone does not improve fatigue performance, even though it can significantly reduce surface roughness. The beneficial effects of the smoother surfaces may have been countered by the volumetric defects close to the surface induced by laser-polishing.

The beneficial effects of the smoother surfaces may have been countered by the volumetric defects close to the surface induced by laser-polishing. The fracture surfaces also reveal that fatigue cracks are initiated from the defects close to the surface, and therefore, fatigue behavior is not improved.Keywords: Inconel 718; Laser powder bed fusion; Laser-polishing; Surface roughness; FatigueIntroductionMetallic additive manufacturing (AM), a layer-by-layer fabrication method, has received significant attention due to the feasibility of building complex geometries such as lattice structures that cannot be constructed by traditional subtractive manufacturing methods [1]. Laser powder bed fusion (L-PBF) is one of the well-known and popular AM methods because it provides arguably the most accurate built dimension and the least surface roughness (SR) through relatively thin layer thicknesses among various AM methods [2].

Although L-PBF fabricates geometrically precise parts, post-processing including thermal and surface treatments are still often applied required to employ L-PBF parts in engineering applications [3].Inconel 718 (IN718) is a well-known nickel-base superalloy in many applications, such as in gas turbines due to its excellent mechanical properties and corrosion resistance [4]. However,474IN718 has high production costs owing to its expensive machinability [5]. Through the near-net- shaping capability of L-PBF, manufacturing of IN718 can achieve reduced machining costs and enable light-weighting. Therefore, IN718 fabricated by L-PBF technique has been studied by numerous researchers to verify practical applicability [6][7][8]. Gruber et al.

Therefore, IN718 fabricated by L-PBF technique has been studied by numerous researchers to verify practical applicability [6][7][8]. Gruber et al. [9] reported that L- PBF IN718 fabricated by optimized process parameters can have comparable hardness and tensile properties to wrought counterparts after appropriate heat treatments including stress relief, hot isostatic pressing (HIP), solution annealing, and standard double aging.Although L-PBF fabricates parts with the smoothest surface among AM methods, the SR is still the most critical factor for fatigue properties. The fatigue lives of as-built (AB) L-PBF parts are significantly less than machined and polished counterparts because SR of AB L-PBF parts can act as micro-notches on the surface resulting in early fatigue crack initiations [11]. Therefore, post surface treatments are still required especially for fatigue critical applications. Lee et al.

Therefore, post surface treatments are still required especially for fatigue critical applications. Lee et al. [12] studied the effects of various post surface treatments on fatigue behavior of L-PBF IN718 and reported that all post surface treatments including sand-blasting, drag-finishing, turning, and grinding can improve fatigue performance compared to the AB condition.The post surface treatments, especially conventional techniques such as milling, turning, and grinding, can significantly improve fatigue performance. However, post surface treatments are not always feasible due to the complex geometries of the parts. In addition, these techniques hinder the benefits of AM which are fabricating complex geometries and near-net-shape parts within a single AM process.

In addition, these techniques hinder the benefits of AM which are fabricating complex geometries and near-net-shape parts within a single AM process. Laser techniques such as laser-polishing (LP) and laser-shock-peening have been employed as non-subtractive post surface treatment techniques for AM parts since they can reconstruct surface topography without wasting materials. Lee et al. [13] showed that the laser- polished L-PBF Ti-6Al-4V not only have better fatigue strength than AB counterparts but also have comparable fatigue lives to machined and polished counterparts.The effects of LP on the fatigue behavior of L-PBF IN718 are examined in this study. Surface topographies and porosities of AB and LP specimens were measured followed by uniaxial strain- controlled fatigue testing to correlate the fatigue properties for each condition. The results show that LP successfully improves the surface quality of L-PBF IN718. However, LP processes generate defects close to the surface.

The results show that LP successfully improves the surface quality of L-PBF IN718. However, LP processes generate defects close to the surface. As a result, LP specimens have similar fatigue lives compared to AB counterparts due to the defects close to the surface that introduce early crack initiations although the surface is smoother.Experimental ProcedureIN718 fatigue specimens were fabricated using an EOS M290 L-PBF system in an argon atmosphere with the default process parameters provided by the manufacturer: laser power of 285W, laser scanning speed of 960 mm/s, hatching space of 110 µm, and layer thickness of 40 µm. According to ASTM E606, with the exception of the gage diameter of 5 mm [14], the geometry of specimens was designed as shown in Figure 1 and the specimens were fabricated vertically. One of the grips has a longer length to provide space to provide enough grip for the LP process.

One of the grips has a longer length to provide space to provide enough grip for the LP process. The specimens were heat treated based on ASTM F3301 and AMS 5564C: stress relief (1065°C for 1.5h followed by furnace cooling), solution treatment (1065°C for 1h, argon quenching), and two-step aging (760°C for 10h, furnace cooling to 650°C and holding for a total precipitation time of 20h) [15,16]. Two different LP conditions (i.e., LP1 and LP2) were processed by two different LP process parameters: LP1 – power of 400W, frequency of 600 Hz, scanning475speed of 100 mm/s, pulse duration of 0.5 ms, defocus of 14 mm and LP2 - power of 250 W, frequency of 700 Hz, scanning speed of 50 mm/s, pulse duration of 0.7 ms, defocus of 14 mm.Figure. 1 Fatigue specimens’ geometry according to ASTM E606.Surfaces of AB, LP1, and LP2 specimens were scanned by a Keyence VHX-6000X optical microscope using 500X magnification with a scanned area of 2 mm by 1 mm.

In order to obtain surface height data, a 3D stitching function was used. The captured 3D images were analyzed by software provided by Keyence. The 3D images were converted to the CSV files to calculate SR values using MATLAB.The porosities of AB, LP1, and LP2 specimens were measured by Zeiss Xradia 620 Versa X-ray computed tomography (XCT) system. The gage sections with a diameter of 5 mm were scanned using the following settings: voltage of 160 kV, current of 156 µA, and voxel size of 5.5 µm. The scanned files are reconstructed by the software provided by Zeiss and saved as TIFF image files. The obtained TIFF image files were post processed by ImageJ and Dragonfly to visualize defects.Fully reversed uniaxial strain-controlled fatigue tests with the constant strain rate of 0.0025 mm/mm/s were conducted on an MTS Landmark hydraulic system according to ASTM E606 [14].

Five different strain amplitudes (ɛa = 0.006, 0.005, 0.004, 0.003, and 0.0025 mm/mm) were tested for all three surface conditions to evaluate their influences in both low and high cycle fatigue regimes. If plastic deformation was not observed from the hysteresis loops during strain-controlled tests, the extensometer was removed after a few thousand cycles and the tests were switched to the force-control to reduce the testing times by increasing frequency.Results and DiscussionThe surface topographies of AB, LP1, and LP2 specimens generated by Keyence software are presented in Figure 2. The calculated SR values such as the arithmetical mean height (Ra), the maximum profile valley depth (Rv), and the maximum height of profile (Rz) are listed in Table 1. The AB specimen has partially melted powder particles and layer-wise undulations on the surface as shown in Figure 2(a).

The AB specimen has partially melted powder particles and layer-wise undulations on the surface as shown in Figure 2(a). Both LP1 and LP2 not only removed partially melted powder particles but also reconstructed surfaces by remelting surfaces. As a result, Ra and Rv, which are considered important SR values in fatigue applications, are reduced by LP processes.476Figure. 2 Surface topographies of AB, LP1, and LP2 specimens overlaid with micrographs and colormaps.Table. 1 The calculated SR values including Ra, Rv, and Rz of AB, LP1, and LP2 specimens.Surface conditionsAB LP1 LP2Ra (µm) 4.75 2.91 2.35Surface roughness parameters Rv (µm) 12.23 8.12 9.16Rz (µm) 39.19 18.32 19.54Visualized XCT 3D images of AB, LP1, and LP2 specimens are presented in Figure 3 and porosity measurement results are listed in Table 2. AB specimens have significantly less defects in both number and size compared to LP1 and LP2 specimens.

AB specimens have significantly less defects in both number and size compared to LP1 and LP2 specimens. Interestingly, they have a similar number of defects in the center but LP1 and LP2 have slightly more defects close to the surface compared to AB, which were likely induced by the LP process itself. According to Table 2, the number of defects for both LPs was almost identical although the process parameters for LP1 and LP2 are different. The correlation might be that the total energy input per unit surface area for both parameters was kept almost the same to each specimen (LP1 ~0.2 J/mm2 and LP ~0.175 J/mm2) and therefore a similar melt pool depth and width due to LP processes might be expected. In order to remedy the LP process induced defects, HIP might be required as a post treatment, especially for fatigue critical applications.477Figure. 3 Visualized XCT 3D images of AB, LP1, and LP2 specimens.Table.

3 Visualized XCT 3D images of AB, LP1, and LP2 specimens.Table. 2 Porosity measurement results of AB, LP1, and LP2 specimens.Surface conditions Number of defects Maximum volume of defects (mm3) Porosity (%)AB 42 3.83E-05 0.0005LP1 632 1.21E-04 0.0074LP2 622 1.17E-04 0.0078Strain amplitude versus reversals to failure fatigue plots of AB, LP1, and LP2 specimens are generated based on fully reversed uniaxial strain-controlled fatigue test results and presented in Figure 4. The effects of SR on fatigue behavior are typically significant in the high cycle regime since SR act as micro-notches that can induce early fatigue crack initiation [11]. However, LP1 and LP2 specimens have similar fatigue lives compared to AB specimens even in the high cycle regime (i.e., at ɛa = 0.003 and 0.002 mm/mm), although SR is decreased.

However, LP1 and LP2 specimens have similar fatigue lives compared to AB specimens even in the high cycle regime (i.e., at ɛa = 0.003 and 0.002 mm/mm), although SR is decreased. Other surface treatments such as sand-blasting and drag-finishing were shown to successfully improve fatigue performance, although SR of them are similar to or even higher than LP1 and LP2 [12]. It indicates that other factors such as defects or tensile residual stresses induced by LP processes can hinder any expected improvements in fatigue strength due to the lower SR.478Figure. 4 Strain amplitude versus reversals to failure fatigue plots of AB, LP1, and LP2 specimens.Fracture surfaces of AB, LP1, and LP2, specimens tested at the lowest strain amplitude (i.e., ɛa = 0.002 mm/mm) are presented in Figure 5. As shown in Figure 5(a), the AB specimen shows multiple crack initiation sites from the surface.

As shown in Figure 5(a), the AB specimen shows multiple crack initiation sites from the surface. LP1 and LP2 specimens not only have multiple crack initiation sites at the surface but also crack initiations from the defects close to the surface as shown in Figures 5(b) and (c). It confirmed that fatigue performances of LP1 and LP2 are not improved partly because of the defects close to the surface produced through LP processes. The two different process parameters utilized for LP processes were both prone to form defects. HIP process and two-step aging followed by LP processes can be suggested to close minimize defects without changing process parameters and improve fatigue strengths.479Figure. 5 Fracture surfaces of AB, LP1, and LP2, specimens tested at ɛa = 0.002 mm/mm (crack initiation sites are shown by the yellow arrows).ConclusionsIn this study, the effects of laser polishing on the fatigue behavior of L-PBF IN718 have been investigated.

Two different laser polished surface conditions and the as-built surface condition were considered. Through surface and porosity measurements, fatigue tests, and fractography analyses, the following conclusions were drawn:The laser-polishing can remove partially melted powder particles and reconstruct surface topography by remelting the surface. The surface roughness of laser-polished specimens is lower than as-built specimens.The laser-polishing produced defects, which resulted in a higher number of defects near the surface in the melted layer. The laser-polishing process parameters such as defocus,480speed, power, and pulse duration should be considered in further study to reduce creating defects.Fatigue life cannot be improved by laser-polishing regardless of the process parameters of laser-polishing processes. Since the tensile residual stresses and defects produced by laser- polishing processes can cause early crack initiations resulting in shorter fatigue lives.

Since the tensile residual stresses and defects produced by laser- polishing processes can cause early crack initiations resulting in shorter fatigue lives. • Hot isostatic pressing is recommended to improve fatigue strengths by reducing defects close to the surface without changing laser-polishing process parameters.AcknowledgmentThis material is based upon work partially supported by the National Science Foundation (NSF) under grant # 1919818 and National Institute of Standards and Technology (NIST) under Award No. 70NANB18H220. The work also was supported from ERDF "Research of additive technologies plus" (No.CZ.02.1.01/0.0/0.0/18_069/0010040).forfutureapplicationsin machineryindustryRTIReferences[1]Frazier WE. Metal additive manufacturing: A review. J Mater Eng Perform 2014;23:1917–28. https://doi.org/10.1007/s11665-014-0958-z.[2] King WE, Anderson AT, Ferencz RM, Hodge NE, Kamath C, Khairallah SA, et al.

e Effect of Layer Orientation on the Tensile Properties of Net Shape Parts Fabricated in StereolithographyAndrew C. Layton and David W. Rosen Manufacturing Research Center Georgia Institute of Technology Atlanta, Georgia 30332Abstract Stereolithographic technologies create parts in thermoset plastic polymeric mixtures of acrylates and epoxies. In order to predict the mechanical behavior of these parts, it is critical to understand the effects that build parameters have on the final properties of the polymer. Using a statistics based approach, the build parameters of layer orientation, layer thickness, and resin class are used as inputs. The response variables, peak stress, elongation at break and Young’s modulus (modulus of elasticity), are examined using the methodology specified in ASTM D638- 01 with modifications as noted.

The response variables, peak stress, elongation at break and Young’s modulus (modulus of elasticity), are examined using the methodology specified in ASTM D638- 01 with modifications as noted. An initial test in Somos 8120 showed the surprising (and statistically significant) result that load bearing capability in the build direction was greater than in the in-layer direction. Additional tensile tests in Somos 8120 and Vantico SL-5510 were undertaken to verify this result, and determine if this effect is present across different classes of resin. This report details the rationale behind this experiment, presents the results to date, and outlines future efforts.1.0 It is reasonable to assume that material property anisotropies exist due to the nature of layer- based fabrication processes.

A sizable body of published work shows material anisotropies in the Fused Deposition Modeling (FDM) [1], and Selective Laser Sintering (SLS) [2][3] technologies; however, there is not a corresponding sizable body of published work with respect to stereolithographic (SLA) processes and photopolymers. Some qualitative discussions of the stereolithography materials in the early 1990’s recognized that material anisotropies were likely; however, the differences should not be statistically significant. There does not seem to be a quantitative work in the public record to verify this supposition, though there may be information in the proprietary domain.IntroductionAs part of a larger effort to characterize SLA materials, a simple screening experiment was performed.

The intent was to verify that the expected material anisotropy based on the relationship of build direction to tensile loading direction was small enough (i.e., statistically insignificant) that we could build tensile test coupons in any convenient orientation within the vat.Stereolithography processes build parts by irradiating a homogeneous photopolymer resin in liquid form. The irradiating energy is provided by an ultra-violet laser. The photoinitiators in the resin system are formulated to react to the specific wavelength of the laser and initiate a polymer chain addition reaction when irradiated. An optic system focuses the laser energy to a discrete spot on the surface of the liquid resin. Laser spot location and motion are controlled289through the mirror and galvanometer set. By vector scanning the laser spot, a line of solidified material is formed. Successive overlapping vector scans create a series of bonded lines to form a layer.

By vector scanning the laser spot, a line of solidified material is formed. Successive overlapping vector scans create a series of bonded lines to form a layer. A second series of vector scans, moving the laser spot in a series of overlapping scans 90o from the previous direction, helps ensure complete reaction of the material, and formation of a good intra-layer bond.A 3-dimensional object is created by bonding layers together. Subsequent layers are created by depositing additional resin on the surface of the previous layer, and repeating the hatching and filling processes. The laser energy solidifies the current layer as described above, but some energy is imparted to the previous layer. This energy causes a slight overcure of the previous layer, and causes the inter-layer bond to develop.The solidified lines that form a single layer are irradiated in an overlapping pattern within a short time frame.

This should have the effect of initiating very high levels of cross-linking between the individual scan lines within a layer. High levels of cross-linking would seem to indicate high strength. The inter-layer bonding is initiated by the ‘left-over’ energy not absorbed by intra-layer cross-link formation. The energy felt at the liquid / solid interface should be more variable than that felt in the liquid. Since the solidified layer has already developed a cross-link network, the number of potential cross-linking sites for inter-layer bond formation should be substantially less than the number available for intra-layer bonding. These two factors should lead to the intuitive conclusion that intra-layer strength is superior to inter-layer strength.2.0 Experimental Method Three sets of experiments were performed in this study: (1) a screening experiment in Somos 8120; (2) a more in-depth experiment in Somos 8120; and (3) a validation experiment in Vantico SL-5510.

For the first two experiments, the test coupons were built in an SLA-3500 using the standard “Fast” build style. The test coupons for the validation experiment were built in an SLA Viper Si2 using the standard “Exact” build style in the low-resolution mode. Both of these SLA machines use Nd:YVO4 laser systems with output wavelengths of 355 nm, and all of the test coupons were built from the same series of .stl files. In all three experiments, testing was within 24 hours of fabrication.2.1 Screening Experiment in Somos 8120. For the initial screening experiment in DSM 8120, three sets of thirty-six samples were tested. (1) Built flat in the XY plane – Flat group; (2) edgewise in the XY plane – Edge group; and (3) standing along the Z-axis – Standing group.

(1) Built flat in the XY plane – Flat group; (2) edgewise in the XY plane – Edge group; and (3) standing along the Z-axis – Standing group. We were not able to gather valid data with respect to the peak stress, elongation at break and modulus of elasticity of the Standing group; however, we were able to gather valid data with respect to ultimate load for all three groups. We chose to use ultimate load as a proxy value for the peak stress in comparing the three groups.Given the use of the ultimate load as a proxy for peak stress, and the surprising result, we felt it prudent to replicate the experiment, gather valid data for peak stress, elongation at break and Young’s modulus, and reexamine the result.

Realizing that the Flat and Edge groups were simply variations on a theme, and recalling that maximum shearing stress occurs on a 45o offset plane [4], we redefined the sample orientation set to include in-plane, transverse, and shear loaded samples.290For this experiment, we considered three orientations 2.2 Second experiment in Somos 8120. designated as Flat (in-plane), Standing (transverse), and Angled (shear). The three orientations are illustrated in Figure #1.ZYAngledStandingFlat45oFigure 1. The three representative sample orientations with respect to the XYZ frameWe arrived at these three orientations by using the following rationale. The primary axes and 45o offsets from the primary axes combine to describe seven (7) unique orientation vectors. If we allow one degree of freedom, a 90o rotation around the vector axis, we create fifteen (15) unique sample orientations.

If we allow one degree of freedom, a 90o rotation around the vector axis, we create fifteen (15) unique sample orientations. We see that these fifteen sample groups represent five distinct inter- layer orientations with respect to the “longitudinal - long transverse - short transverse” frame of the individual samples. Each of these five inter-layer orientations is comprised of three sample groups.Within these five orientation plane groups, there are three primary groups – “Long-LT” and “Long-ST” orientation planes are both “In-plane” with respect to the loading direction, both “A” and “B” orientation planes are “shear,” and the “LT-ST” orientation plane is “transverse.” Figure #2 provides a graphical representation of the five distinct inter-layer orientations and their subsequent grouping into the three primary groups denoted.

For the validation experiment in Vantico SL- 5510, we fabricated 12 test coupons for each of the same primary groups tested in the second Somos 8120 experiment. Post-processing, curing, and testing were consistent with the standard methods and uniform across each experiment.2.4 Methods and Apparatus. While exposure to background UV, relative humidity and heat have been shown to affect the properties of parts fabricated in the stereolithography process, we have not sought to control them in this study. The broad spectrum of users neither actively nor consistently controls these factors. One of the purposes of this research is to examine what the user would see outside of the standard conditions. The ASTM standard is designed to provide an “apples-to-apples” comparison of materials – control as many factors as possible and allow the response variable to be driven overwhelmingly by the factor of interest.

This is a valid scenario, but it does not necessarily represent what the user in the field will experience.Similarly, the ASTM D638 standard prescribes fabrication of the test coupons by either die cutting or machining from flat material (e.g., sheet), or by molding. Neither of these matches the SLA process particularly well – generally, SLA parts are made to net shape and used. There may be some surface finishing or treatment, but mostly the surfaces are used as fabricated. By fabricating the test coupons to net shape, we leave the edge effects extant. So, we have chosen to examine those factors (layer orientation, layer thickness, and resin selection) that the user actively controls.

So, we have chosen to examine those factors (layer orientation, layer thickness, and resin selection) that the user actively controls. For the purposes of this experiment, we do conform to the specifications found in the ASTM standard (coupon form and dimensions, strain rates, etc) not discussed above.292We have discussed layer orientation in the section above, the other parameters of interest are:Layer thickness (Quantitative) – This is a discrete parameter that can be set at one of three values - .002 inches, .004 inches, or .006 inches.Resin Class (Qualitative) – This refers to the general class of materials represented by each of the subject resins.Layer thickness is a user selectable parameter based on the build style chosen. Advanced users have the capability to create and tailor custom build styles. Like vat orientations, the potential number of different styles that can be created is unlimited.

Advanced users have the capability to create and tailor custom build styles. Like vat orientations, the potential number of different styles that can be created is unlimited. For the purposes of approximating typical user results, we limit our build styles and, consequently, layer thickness to those provided by the resin manufacturers.Resin Classes - There are three classes of material in the stereolithography process: polyethylene-like / polypropylene-like (PE/PP); general purpose; and ABS-like. Each has a different combination of rigidity and durability properties. Parts made from PE/PP - like resins have low rigidity and high durability while GP resins yield high rigidity and low durability. ABS-like materials are highly rigid and highly durable.The chemical composition of the resins within classes is similar, so a commonly used resin should provide a good representation of its class; however, there are differences in chemistry between the classes.

Differences in the cross-linking reaction are a potential cause for differences in the material properties. By using a representative resin from each class, we seek to gain data that will help us define this phenomenon as being process related, chemistry related, or an interaction of process and chemistry. For our purposes we will use these resins to represent the three classes.General purpose (Epoxy Acrylate): Polyethylene-like: ABS-like:Vantico SL-5510 DSM Somos 8120 Vantico SL-7560The particular formulations and compositions of the resin systems are proprietary. A review of the constituents of the resin systems shows, even without a precise delineation of constituents, that the materials are dissimilar enough in nature to allow us to declare them to be different [5], [6]. This is borne out by an examination of the experimental data values.

This is borne out by an examination of the experimental data values. The differences in the measured loads, stresses and elongation are large enough to see by observation without need for comparison of the mean values by statistical methods. After fabricating the test coupons and post-processing them according to standard practice, we subjected the coupons to tensile testing in an Instron Tensile Tester model 4466. The standard manufacturer provided software (Instron Series IX Automated Material Tester – version 8.11.00) was used for reporting of the pertinent data.293Figure 3: Typical test coupons before and afterFigure 4: Instron tensile testing set-up2.5 Statistical methods. The tensile testing data was subjected to Analysis of Variance (ANOVA) for a fixed effects model per standard and accepted methodologies [7]. As a short review - ANOVA allows us to partition the total variability of the data into its component parts.

As a short review - ANOVA allows us to partition the total variability of the data into its component parts. In our experiment, these components are: (1) the variability between groups due to differences in treatment, and (2) variability within a particular group due to random error. Dividing the components by the degrees of freedom, we have an estimated sample variance of each component. The estimate of the sample variance between treatment groups is denoted MSt, and the estimate of the sample variance within groups due to random error is denoted MSe.Applying Cochran’s theorem, we compare the two components. If the null hypothesis, (that the treatments are equal) is true, then the estimated sample variances should both be independently distributed Chi-square random variables. The ratio of the two components is distributed as an “F” statistic.

The ratio of the two components is distributed as an “F” statistic. Standard “F” test tables give a reject value based on the level of confidence, the degrees of freedom of the treatments, and the degrees of freedom of the random errors.Our test statistic is derived by taking the ratio MSt / MSe. Any value of this test statistic greater than the standard table value allows us to reject the null hypothesis. If we cannot reject the null hypothesis, then performing a contrast gives us no additional information. We have chosen a 1% level of confidence - meaning that the probability of falsely rejecting the null hypothesis as a result of this comparison to the standard table values is less than 1%.Since we are testing more than two distinct groups within each material property and class, a rejection of the null hypothesis would require further analysis contrasting each possible pair of treatment groups.

This gives us additional information regarding the sources of the differences between the individual groups, and relationships between the groups.294N k nid a o L3.0 Experimental Results 3.1 Screening Experiment. For the initial screening experiment, an analysis of variance of the ultimate load data associated to Figure #5 showed that we were able to reject the initial hypothesis - that build direction anisotropies were not significant in load carrying capability - with a much less than 1% chance of falsely rejecting. The test statistic value needed for the 1% chance of false rejection was 4.98 (or greater). Our analysis of experimental data yielded a test statistic of 57.81 - an order of magnitude greater than the “reject” value.Contrasts of the Flat, Edge and Standing data groups showed a statistically significant difference between the Flat and Standing groups, and between the Edge and Standing groups.

The Flat and Edge groups did not demonstrate a statistically significant difference. These contrasts make sense.Build Orientation Effects700650600550Highest Value500Upper QuartileStandingMedianEdgewise450Lower Quartile400350Lowest ValueFlat3000123OrientationFigure 5: Load carrying capability data for the initial experimentThe difference in means between the Flat and Edge groups is less than 5%, and all of the data values of the Edge group fit between the high and low extremes of the Flat group. On further examination, one can see that both the Flat and Edge groups experience in-plane loading during tensile testing. This is not where one would anticipate the anisotropy to be evident. Rather this is an expected result of comparing two variations of in-plane loading samples.2954The difference in mean of the Standing group with respect to the means of the Flat group and the Edge group is +25% and +20%.

It is not surprising to see differences between the mean of the transverse loaded Standing group and the means of the in-plane loaded Flat and Edge groups. This is where one would expect to see evidence of an anisotropy; however, it is surprising that the mean of the Standing group exceeds the means of the Flat and Edge groups by fairly large, and statistically significant, percentages.3.2 The Second Experiment in DSM 8120 material yielded a similar result with additional information. In this experiment, we examined three hypotheses based on the loading direction versus build layer orientation relationship, namely this relationship would not cause statistically significant anisotropies in Peak Stress, Young’s Modulus, and Elongation at Break for test coupons aged less than 24 hours. Also, when an anisotropy was identified, a contrast of the data groups was undertaken.

Also, when an anisotropy was identified, a contrast of the data groups was undertaken. A summary of the test data and contrast results is provided in Figure #6.Different fromDifferent fromDifferent fromParameterGroupMeanStd DevFlatStandingAngledFlat2098.3563.15--YesYesPeak Stress (psi)Standing2495.53131.47Yes--YesAngled2238.1443.20YesYes--Flat10.320.39--YesYesYoung's modulus (ksi)Standing12.900.52Yes--YesAngled11.130.41YesYes--Flat10.190.78--NoYesElongation (%)Standing9.051.57No--NoAngled8.180.66YesNo--Figure 6: Summary of 8120 data for the second experimentPeak Stress - The analysis of variance for the peak stress data showed that we were able to reject the initial hypothesis - that build direction anisotropies were not significant in peak stress - with a much less than 1% chance of falsely rejecting.

The test statistic value needed for the 1% chance of false rejection was 5.39 (or greater). Our analysis yielded a test statistic of 24.96 - an order of magnitude greater than the “reject” value.A statistical contrast of the two peak stress data groups with the least difference in means, the Flat and Angled groups, showed statistically significant differences between those groups. Contrast bears out a statistically significant difference between the Flat and Standing groups as well. The Angled and Standing groups were contrasted as well. These two also proved to have statistically significant differences.

The Angled and Standing groups were contrasted as well. These two also proved to have statistically significant differences. In all of the contrast cases, the test statistic values were an order of magnitude greater than the required rejection value, and each of the groups was different from the others with respect to the Peak Stress property.Young’s modulus - An analysis of variance for the Young’s modulus data showed that we were able to reject the initial hypothesis - that build direction anisotropies were not significant in296Young’s modulus - with a much less than 1% chance of falsely rejecting. The test statistic value needed for rejection was 5.39 (or greater). Our analysis of the experimental data yielded a test statistic of 38.78 – again, an order of magnitude greater than the “reject” value.Contrasting the groups Flat and Angled, we see a statistically significant difference in the Young’s modulus.

Subsequent contrast of the remaining two possible comparisons shows that they are different as well with respect to the Young’s modulus property.Elongation at Break – The initial analysis of variance of the entire data set of three groups shows that we can reject the hypothesis - that build direction anisotropies were not significant in Elongation at Break – with a less than 1% chance of falsely rejecting that premise. The test statistic value needed for the 1% chance of false rejection was 5.39 (or greater). Our analysis of experimental data yielded a test statistic of 10.45.In this case, the Flat and Angled groups showed the largest difference in the group means. The least difference in means was between the Standing and Angled groups. The test statistic for the Standing and Angled contrast was 3.09 versus a minimum reject value of 7.95 – we cannot reject the hypothesis based on differences in these two groups.

The test statistic for the Standing and Angled contrast was 3.09 versus a minimum reject value of 7.95 – we cannot reject the hypothesis based on differences in these two groups. A similar result occurs with the Standing and Flat groups – the test statistic is 5.14 versus a reject value of 7.95 or greater. Only in the contrast of the groups with the largest difference in means, groups Flat and Angled, do we have a test statistic sufficiently large (46.66) to reject the null hypothesis.3.3 Validation Experiment in Vantico SL-5510 - As noted in the Experimental Method section, the suspected material property anisotropies may be related to the chemical composition of a particular class rather than being process related, or evident across all classes. A similar series of tests were run using net-shaped coupons fabricated from Vantico SL-5510 resin.

A similar result in this material may give an indication that the effect is process related.Different fromDifferent fromDifferent fromParameterGroupMeanStd DevFlatStanding AngledFlat5762.32152.26--YesYesPeak Stress (psi)Standing7755.67972.35Yes--NoAngled7837.70838.50YesNo--Flat51.233.43--NoYesYoung's modulus (Ksi) Standing55.769.70No--NoAngled59.338.26YesNo--Flat33.903.70--YesYesElongation(%)Standing21.706.40Yes--NoAngled22.307.40YesNo--Figure 7: Summary of test data and contrast results in Vantico SL- 5510 resin297Peak Stress - The analysis of variance for the peak stress data showed that we were able to reject the initial hypothesis - that build direction anisotropies were not significant in peak stress - with a much less than 1% chance of falsely rejecting. The test statistic value needed for the 1% chance of false rejection was 5.39 (or greater).

The test statistic value needed for the 1% chance of false rejection was 5.39 (or greater). Our analysis of yielded a test statistic three orders of magnitude greater than the “reject” value.A statistical contrast of the two peak stress data groups with the least difference in means, the Standing and Angled groups, showed no statistically significant differences between those groups. The test statistic for these two groups was .05 versus a reject value of 7.95 or greater, and the hypothesis that the two treatment means are different cannot be rejected. Contrasting the two remaining sets leads to the rejection of the hypotheses in both cases.

Contrasting the two remaining sets leads to the rejection of the hypotheses in both cases. The Standing and Flat contrast and the Angled and Flat contrast reject the hypothesis with test statistics of 49.22 and 71.17, respectively.Young’s modulus - Similarly, an analysis of variance for the Young’s modulus data showed that we were able to reject the initial hypothesis - that build direction anisotropies were not significant in Young’s modulus - with a much less than 1% chance of falsely rejecting. The test statistic value needed for the 1% chance of false rejection was 5.39 (or greater). Again, the analysis yielded a test statistic three orders of magnitude greater than the “reject” value.Group contrast results are summarized in Figure #7.

Again, the analysis yielded a test statistic three orders of magnitude greater than the “reject” value.Group contrast results are summarized in Figure #7. Even though the ANOVA of the full data set rejects the hypothesis with a test statistic much larger than the required reject value, we see that two of the contrasts do not show a statistically significant difference between treatment means of the groups under consideration. Neither the Angled and Standing contrast nor the Flat and Standing contrast is capable of rejecting the hypothesis. In these two contrasts, the test statistics were .94 and 2.33, respectively while the reject value was 7.95 or greater.Elongation at Break – The initial analysis of variance of the entire data set of three groups shows that we can reject the hypothesis - that build direction anisotropies were not significant in Elongation at Break – with a less than 1% chance of falsely rejecting that premise.

The test statistic value needed for the 1% chance of false rejection was 5.39 (or greater). Our analysis of experimental data yielded a test statistic of 15.75 – nearly a factor of 3 greater than the “reject” value.In this case, both the Flat and Angled group contrast and the Flat and Standing group contrast generated test statistics sufficiently large to reject the hypothesis. The test statistics generated were 24.06 and 32.96, respectively, versus reject statistic of 7.95 or greater. The test statistic for the Standing and Angled contrast was .04 versus a minimum reject value of 7.95 – we cannot reject the hypothesis.Summary of results – Statistically significant material anisotropies related to build orientation were shown for the Peak Stress, Young’s modulus and Elongation at Break properties of net shaped test coupons fabricated using the Stereolithography process.

This anisotropic effect was seen in two chemically different classes of material.2984.0 Conclusions The initial screening experiment yielded two results: First, that there is a statistically significant anisotropy in the ultimate load carrying capability as a result of the relationship between build orientation and loading direction; and secondly, that the intuitive solution that the in-plane (Flat) property would be superior to the transverse plane (Standing) property was found not to hold in this case. The first result may be expected, but the second result is somewhat of a surprise.The second experiment confirmed the finding of the first experiment and amplified the finding by yielding data on the Peak Stress, Young’s modulus, and Elongation at Break properties. The in-plane loaded (Flat) coupons showed the lowest Peak Stress and Young’s modulus and exhibited the greatest Elongation at Break.

The in-plane loaded (Flat) coupons showed the lowest Peak Stress and Young’s modulus and exhibited the greatest Elongation at Break. The transverse loaded (Standing) coupons showed the highest Peak Stress and Young’s modulus while the least Elongation at break was found in the shear loaded (Angled) orientation.The third experiment set had a similar finding – statistically significant anisotropies in Peak Stress, Young’s modulus, and Elongation at Break – in a different class of material. Again the in-plane loaded (Flat) coupons showed the lowest Peak Stress and Young’s modulus and exhibited the greatest Elongation at Break; however, in this trial, the transverse loaded (Standing) and shear loaded (Angled) coupons switched rankings.

The shear loaded (Angled) coupons showed the highest Peak Stress and Young’s modulus while the least Elongation at break was found in the transverse (Standing) loaded orientation.The similar results with regard to the in-plane loaded coupons (Flat) across all three trials may be an indication of a process driven effect. The switching of ranks of the transverse loaded (Standing) and shear loaded (Angled) coupons in the results of the second and third experiments may be an indication of some process-chemistry interaction. The only two orientations that were found to be statistically different in every contrast, regardless of material, were the shear loaded (Angled) and in-plane loaded (Flat) orientations.These results support the existence of a process driven effect with respect to tensile properties as a function of build orientation.

Also, the data suggests that there may be a secondary process- chemistry interaction.5.0 Future work The next series of experiments will seek to verify the process driven effect by testing the third class of material. As noted, we will be using Vantico 7560 to represent the ABS-like class of materials. Other materials run on equipment other than our own, my yield additional information. In addition, to gain further information on the process driven effect, we will be testing coupons built in 7110 on a HeCd laser system. Given the proprietary nature of the materials, verification of a process-chemistry interaction could be quite difficult to verify.Also, we will be checking this effect at various stages of material aging to determine if it is consistent across time, or simply transient.2996.0 Acknowledgments The RPMI is supported by member companies.

We would like to thank 3D Systems, Baxter International, DSM Somos, Pratt & Whitney, Siemens, and Vantico for their long-standing commitment and generous support.7.0 References [1] Montero, M; Roundy, S; Odell, D; Ahn, S-H; and Wright, P. K., “Material Characterization of Fused Deposition Modeling (FDM) ABS by Designed Experiments,” Society of Manufacturing Engineers, 2001.[2] Agarwala, M. K.; Bourell, D; Beaman, J. J., “Densification of Selective Laser Sintered Metal Parts by Hot Isostatic Pressing,” Solid Freeform Fabrication Proceedings, 1994, pp 65-73. (Discusses ansiotropy of linear shrinkage.)[3] Subramanian, P. K.; Vail, N. K.; Barlow, J. W.; and Marcus, H. L., “Anisotropy in Alumina processed by SLS,” Solid Freeform Fabrication Proceedings, 1994, pp 330-338.

This work shows that recent developments in materials have increased the suitability of DMLS tools for injection moulding processes, in terms of the force required to eject a part from a tool.Tools were manufactured for testing from two bronze (DirectMetal 50 and DirectMetal 20) and two steel (DirectSteel 50 and DirectSteel 20) powders, and the forces required to eject ABS parts from these tools were recorded. It can be seen that the more recent powder developments, sintered in 20 µm layers, showed a dramatic improvement over the older powders, sintered in 50 µm layers.1.0IntroductionThe benefits of using Rapid Prototyping methods to create tooling have been shown throughout industry, in terms of the both the time and the cost required to produce the tooling.

In some cases a reduction of 50% in each has been achieved by using Rapid Tooling techniques rather than conventional methods (Lohner, 1996).Other well-publicised benefits of Rapid Tooling are the ease of producing intricate and complex geometries (Nelson, 2000), and the ability to produce thin-walled sections, which are generally difficult to machine (Dvorak, 1993). Internal geometries that would previously have been impossible to manufacture, can now be produced with relative ease. This allows the possibility of incorporating conformal channels within a tool, to provide better heating or cooling to specific areas of importance (Himmer et al, 1999, Dalgarno et al, 2001).However, as with any technology, there are also some disadvantages to be found. The ‘stair-stepping’ effect inherent to all Rapid Prototyping processes can cause difficulties when attempting to remove a moulded part from a tool.

The ‘stair-stepping’ effect inherent to all Rapid Prototyping processes can cause difficulties when attempting to remove a moulded part from a tool. In the case of laser-sintered tools, their porosity can further inhibit part removal, as moulded material ingresses into the tool itself.73High forces required to remove a part from a mould can lead to warpage and distortion of the parts, or in some cases parts becoming completely stuck. Another area of concern is in tools with fine or delicate features which are simply not strong enough to withstand the forces required. These features can become distorted or even break off completely.

These features can become distorted or even break off completely. Material and equipment developments are therefore necessary in order to improve the surface quality of tools made using Rapid Tooling techniques, and consequently to reduce the ejection forces required.2.0Development of DMLSThe early stages of the development of direct metal laser sintering (DMLS), as with mostRapid Prototyping processes, focussed mainly on the speed of the process, rather than on the properties of the sintered part (Lind et al, 2002).The very first powder developed for use in DMLS was a bronze powder to be sintered in 100 µm layers. Although this provided the ability to build parts relatively quickly, the amount of finishing work required to give the parts an acceptable surface finish could negate this time- saving.

Although this provided the ability to build parts relatively quickly, the amount of finishing work required to give the parts an acceptable surface finish could negate this time- saving. After the development of this early material, steel and bronze based powders were developed in 50µm layers, and most recently both powders have been developed in 20µm layers.The smaller steps encountered on a Rapid Prototyped tools built with a small layer thickness, and in a finer powder, provide a naturally better surface finish than when built with a larger layer thickness. When using DMLS some finishing will generally be necessary once the build is complete, particularly if the part is to be used as a tool for injection moulding. Whereas previously large amounts of finishing would be required in order to render the product usable, shot peening is now commonly used to finish DMLS parts (Lind et al, 2002).

Whereas previously large amounts of finishing would be required in order to render the product usable, shot peening is now commonly used to finish DMLS parts (Lind et al, 2002). Injection moulding tools made from the newer powders have been shown to require only shot-peening to give similar performance to an EDM tool in terms of ejection force (Majewski, 2001).3.0Research AimThe newer powder developments for the DMLS process can be expected to give improvedsurface quality on tools produced in this way. The work described here quantitatively evaluates this improvement in terms of the ejection forces required for parts injection moulded into DMLS tools.4.0 MethodologyTools were produced with varying levels of finishing, in each of the four powders now available for use in DMLS – DirectMetal 50 and DirectMetal 20 (bronze-based powders) and DirectSteel 50 and DirectSteel 20 (steel-based powders).

The following sections describe the tool design and moulding parameters used.744.1Tool DesignPrevious work (Hopkinson 2000) on stereolithography tools involved a simple tooldesign; the same geometry has been used in this work. The design of the tool was intended to provide a good contact area with the moulded part, but also to minimize the time spent on sintering. The cavity side of each tool was therefore machined, as it would have no effect on the force required to eject the part. Another benefit of using a machined cavity was to reduce the possibility of the moulded part from being retained in the cavity side of the tool on mould opening.The core side of the tool consisted of a machined surround to be used for each tool, withindividual inserts screwed into this surround. In order to provide easier ejection, most inserts were shot-peened to give a lower surface roughness, and a draft angle of 0.5°.

In order to provide easier ejection, most inserts were shot-peened to give a lower surface roughness, and a draft angle of 0.5°. Figure 1 shows the tool design and geometry used.Machined core surroundRemovable coreMachined cavity surroundTest inserts produced by different routesMouldingFigure 1 - Tool and part geometryWhilst a simple geometry was chosen (particularly to enable easy production of the machined inserts), the inserts had a long line of draw. This provided a large area of contact between the mould tool and the moulded part, providing adequately difficult ejection to perform the required tests.4.2Injection ParametersThroughout each test, the injection parameters (including pressure, speed andtemperature) were maintained constant. Cooling time and mould opening were also maintained constant.

Cooling time and mould opening were also maintained constant. Table 1 shows the parameters used.75ParameterValueTool Temperature Prior to Injection Injection Temperature Cooling Time Prior to Mould Opening Mould Open Time Prior to Ejection Total Cycle Time40 °C – 50 °C 225 °C 15 seconds 0 seconds ~ 120 secondsTable 1 – Injection parameters4.3 Moulding MaterialThe moulding material was ABS, as it provides a difficult material to mould and eject,and is also widely used in industrial applications of injection moulding. Ejection force measurements were only taken where the part produced was of a good quality, and suffered no distortion or flashing.4.4Number of Parts MouldedIn order to ensure that enough results were taken to give accurate measurements, 20 partswere moulded from each insert. It can be seen from the results section of this report that no substantial variations are apparent over the course of these 20 shots.

It can be seen from the results section of this report that no substantial variations are apparent over the course of these 20 shots. It was therefore considered unnecessary to test a higher number of mouldings.4.5Tool Temperature Prior to InjectionWhilst in some inserts the temperature of the mould did not appear to make a large difference, on certain tools it was noticeable that if the insert was not cooled after each shot then the ejection force measured on the next shot was substantially lower. For this reason the temperature of the mould was maintained between 40° and 50° at the beginning of each moulding cycle.4.6 Measurement of Ejection Force (Fe)The tool design used three ejector pins arranged around the central upstand. Load cells were positioned behind each of these pins, and individual readings taken upon ejection of each part (see figure 2).

Load cells were positioned behind each of these pins, and individual readings taken upon ejection of each part (see figure 2). The individual results taken from each load cell were added together to give a value for the total force required to eject the part. In order to give an accurate reading of the peak force, a sampling rate of 1200Hz was used over a three second ejection period.76Ejector Back PlateLoad CellMoulded PartEjector PinEjector Front PlateFigure 2 – Location of Load Cells4.7Finishing proceduresTable 2 shows the different finishing procedures applied to DMLS tools for this work.Type of Finishing UnfinishedDescription No finishing procedures appliedTime Taken for Finishing N/AShot Peening (Type 1)1. Organic medium to remove loose powder from tool surface10-15 minutes2. Steel medium tosmooth and compact surface3. Organic medium for further smoothingShot Peening (Type 2)1. Steel Medium 2.

Steel medium tosmooth and compact surface3. Organic medium for further smoothingShot Peening (Type 2)1. Steel Medium 2. Ceramic Compound5-10 minutesTable 2 - Description of Finishing Methods5.0ResultsThese tests were designed to compare the performance of injection moulding tools made by DMLS, and to identify any improvement to be found in the newer 20 µm powders. Ejection force measurements were recorded for each of the 20 shots produced on each insert. An insert produced by electro-discharge machining (EDM) was also tested for benchmarking purposes; this is a standard process for producing tools in industry, and is included in this work as an indication of an industrial acceptable level of ejection force. Table 3 below shows the average of these ejection forces.

Figure 3 below shows the ejection forces for each of the steel inserts tested, as well as those for the EDM insert.900800700)N600( e c r o F n o i t c e j E50040030020010001234567891011121314151617181920Shot numberDS20 UnfinishedDS20 Peened (1)DS20 Peened (2)DS50 Peened (1)DS50 Peened (2)EDMFigure 3 - Ejection Forces for Steel Based Powders78It can be seen from these results that neither of the DirectSteel 50 inserts tested, compare favourably in terms of ejection force with the EDM insert tested. The average values recorded (505N for an insert finished with type 1 peening, and 560N for that finished using type 2), are approximately three times higher than that for an EDM insert (170N).In the case of the simple geometry chosen here, the higher ejection forces recorded on these inserts showed no apparent adverse effects on either the part or the tool.

However, when moulding more difficult geometries, perhaps with fine features or thin-walled sections, these higher forces could potentially cause part distortion or damage upon ejection, as well as damaging the tool itself.In the case of the two peened DirectSteel 20 inserts, the average ejection forces, (153N for type 1 and 116N for type 2), are both below that for the EDM tool. This would suggest that these inserts are at least as acceptable for injection moulding as tools made by the EDM process.Also of interest is the average ejection force of 756N recorded on the unfinished DirectSteel 20 insert. Whilst this is still substantially higher than for the EDM insert, it is only approximately 200N higher than that of a DirectSteel 50 peened insert.Shot peening is a standard recommended method of finishing for DMLS injectionmoulding tools, even for those made in the 50µm powders.

On simple geometries therefore, especially where the surface finish of the part is not critical, these results would suggest that even a DirectSteel 20 insert with no finishing may be acceptable as an injection moulding tool. The obvious benefit of this would be a reduction in the overall lead time of a tool, as even peening procedures will add time to the process.It can be seen from these results that there is a marked improvement in the performance of the 20µm powders over the 50µm ones.

Table 4 shows the percentage reduction in ejection force found with the steel based powders.Original MaterialDS50DS50DS50New materialDS20DS20DS20Level of FinishingPeened (Type 1) Peened (Type 2) Average% Reduction in Required Fe 70%79%74.5%Table 4 - Percentage reduction in ejection force (steel based powders)This reduction in ejection forces would be expected to greatly improve the suitability of steel DMLS tools for injection moulding, particularly for more difficult or complex geometries, although it should be noted that a smaller layer thickness will require a greater sintering time.795.2Bronze based powdersThe following section discusses the results gained from the bronze based powders tested –DirectMetal 50 and DirectMetal 20. Figure 4 shows the ejection forces recorded on each shot.

Figure 4 shows the ejection forces recorded on each shot. As before the results for the EDM insert are included as a benchmark.500450400350)N( e c r o F n o i t c e j E3002502001501005001234567891011121314151617181920Shot N umberDM50 Peened (1)DM50 Peened (2)DM20 Peened (1)DM20 Peened (2)EDMFigure 4 - Ejection Forces for Bronze Based PowdersThe two peened DirectMetal 50 inserts recorded very similar average ejection forces of 426N and 437N for type 1 and 2 respectively.

As with the DirectSteel 50 inserts, these forces would not necessarily render tools made in this way unsuitable for injection moulding, but a lower ejection force would in general be preferable.In this case the DirectMetal 50 inserts recorded ejection forces of approximately 2.5 timesthose for the DirectMetal 20 or EDM inserts.With type 1 peening, the average ejection force recorded from a DirectMetal 20 insert was 199N, and for type 2 peening it was 153N. These figures compare extremely favourably with the EDM insert tested, suggesting that they too are extremely suitable for injection moulding tools.As in the case of the steel based inserts, the 20µm bronze powder has shown a markedimprovement over the 50µm.

Table 5 shows the percentage improvements seen here.80Original MaterialDM50DM50DM50New materialDM20DM20DM20Level of FinishingPeened (Type 1) Peened (Type 2) Average% Reduction in Required Fe 53%65%59%Table 5 - Percentage Improvement in ejection force (bronze based powders)The average reduction in Fe of 59% is fairly dramatic, suggesting increased suitability of the 20µm tools for injection moulding. Whilst the percentage improvement seen is less than that for the steel based powders, it can be seen from table 3 that the ejection forces for the 50µm bronze powder were also lower than for the 50µm steel powder. The actual values recorded for all the peened 20µm inserts did not show a large amount of variation.6.0ConclusionsPoor surface finish is one of the most common concerns held by industry over the use ofRapid Prototyped tooling.

However, the recent powder developments for the DMLS process have produced a marked improvement in terms of the ejection force required for injection- moulded parts.In many cases the ejection forces found when using a peened tool made from either of the 50µm powders will be low enough to cause no problems when ejecting injection moulded parts. However, in the case of fragile sections or complex geometries, it is possible that even this level of ejection force may be enough to damage the part and/or the tool.The newer 20mm powders showed dramatic reductions in ejection force of 74.5% and59% for the steel and bronze based powders respectively.

This reduction to the same level as an EDM tool suggests that the surface finish is no longer an obstacle to effective de-moulding from a DMLS tool.It is important to realise that tools sintered in smaller layers will require a longer sinteringtime than with thicker layers, although the surface finish of parts moulded using these tools will generally be better. The choice of powder is then affected by which is the more important of these considerations.It is still the case that a DMLS or Rapid Prototyped tool will not always be the best choicefor injection moulding or any other process. Decisions must be made as to the required surface finish of the part, as well as the acceptable lead-time and cost of the tools.

However, the new developments in materials for DMLS render this process much more suitable for tooling purposes than in the early stages of their development.817.0 ReferencesDalgarno K, Stewart T, Allport, J, Layer manufactured production tooling incorporating conformal heating channels for transfer moulding of elastomer compounds, Plastics, Rubber and Composites, v 30, n 8, 2001, pp 384-388Dvorak P, Rapid Prototypes for Production Parts, Machine Design, 1993, pp48-54Hopkinson N, Dickens P, Using Stereolithography Tools for Injection Moulding: Research into Tensile Tool Failure and Unexpected Benefits of the Process, Proc Institution Mechanical Engineers-B, 2000, 214(B10), 891-89Himmer T, Nakagawa T, Anzai M, Lamination of Metal Sheets, Computers in Industry 39, 1999, pp 27-33Lind J, Hanninen J, Kotila J, Nyrhila O, Syvanen T, Direct Metal Laser Sintering – From 100 µm to 20µm Layers, Proceedings of the 2002 International Conference on Metal Powder Deposition for Rapid Manufacturing, 2002, pp 225 - 232Lohner A, Injection Mould Making the Quick Way – Laser Sintered Cavities and Mould Inserts, Kunstoffe Plast Europe, November 1996, pp 7-9Majewski C, The Effect of Draft Angle and Level of Finishing on the Ejection Force Required when Injection Moulding Using Direct Metal Laser Sintered Tools, M.Sc.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperThe Effect of Location on the Structure and Mechanical Properties of Selective Laser Melted 316L Stainless SteelE. Fitzgerald1*, W. Everhart1 1 National Security Campus, Kansas City, MO 64147The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.AbstractAdditive Manufacturing (AM) has significantly increased the design freedom available for metal parts and provides flexibility within each build to produce multiple components of varying size and shape.

In order to obtain the highest build efficiency, it is ideal to print multiple parts together spanning the entire plate, with as little spacing as possible between the parts. Work has been performed to characterize the variance of materials properties as a function of location within the build volume. This work utilizes mechanical, chemical, and structural analysis techniques to expand on previous work by statistically evaluating the impact of build location on tensile performance in Selective Laser Melting (SLM). Results show that properties span a small range regardless of build design. Specific effects of cover gas convection, location, and build length are observed and evaluated.

Results show that properties span a small range regardless of build design. Specific effects of cover gas convection, location, and build length are observed and evaluated. Finally, sub-grain misorientation values are correlated to material strength through Electron Back Scatter Diffraction (EBSD).IntroductionThe adoption of metal additive manufacturing (AM) has grown over the past decade due in no small part to the newly available geometric freedom and the capability to customize each build based on the number and size of any variety of components. As these new technologies are adapted and used to manufacture goods on larger scales, it is important to understand the variables in the machine and the impact those variables have on the built material. Many studies have investigated the effect of scanning parameters on residual stress, bulk properties and microstructure in AM [1-12].

Many studies have investigated the effect of scanning parameters on residual stress, bulk properties and microstructure in AM [1-12]. These studies include evaluation of the structure, residual stresses, and properties generated by the fast cooling rates present in AM. Secondary build variables, used here to describe variables outside of machine scanning parameters like location or orientation of a part in the build, proximity of neighboring parts and machine operations, may affect local temperature gradients but are commonly assumed to be insignificant.Electron Beam Melting (EBM) is an example of a technology wherein several studies have included information on the impact of secondary build variables on material microstructure and performance [1-4]. N. Hrabe and T. Quinn [1, 2] investigated the location of the part in the build and found no significant effect on the microstructure and material properties.

N. Hrabe and T. Quinn [1, 2] investigated the location of the part in the build and found no significant effect on the microstructure and material properties. They did observe a small statistical difference in yield and tensile strength by part size, but could not determine a relationship between strength and size of the part. In the study samples of each size were printed©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA-0002839.574as a group in different regions of the build chamber, therefore location may play a role, not just size. Additional research into the microstructure throughout the height in a EBM build has shown there was no effect on material density but variability in microstructure can be seen [3,4].

Additional research into the microstructure throughout the height in a EBM build has shown there was no effect on material density but variability in microstructure can be seen [3,4]. This change in microstructure corresponded to a reduction in material strength and hardness.As for Selective Laser Melting (SLM) technology, secondary build variables have been less investigated. Abd-Elghany and Bourell [5] considered a number of variables to control the quality of SLM printed parts. Their work focused on the scanning and layer deposition process, concluding that these factors have an effect on the density and surface finish of printed parts. They found optimization of the processing parameters can result in relatively high density parts. K. Saeidi et al. [6] built on studies of process parameter optimization to investigate the solidification process in SLM printed 316L stainless steel.

K. Saeidi et al. [6] built on studies of process parameter optimization to investigate the solidification process in SLM printed 316L stainless steel. They found the cooling conditions in laser melting were enough to form columnar structures, with sub-grains much smaller than those found in other rapidly cooled manufacturing processes. Scanning transmission electron microscopy (STEM) imaging of the samples indicated the SLM parts had inhomogeneous diffusion in the sub-grain boundaries and chromium silicate nano-inclusions. These features and the high density of the part resulted in high microhardness. Kruth et al. [7] investigated deformation in builds based on powder layer density and scanning strategies, concluding shorter scan length, effective powder coverage and melt of the first layer reduced thermal stresses in the part.

This work highlights the impact of factors outside of typical beam scanning machine parameters, but does not proceed to connect the observed effect these factors have on melt and cooling with microstructural observations.In order to understand SLM as a technology and the structure and performance of the parts it builds, the SLM process and machine factors must be considered as a whole. As the EBM research has shown, there is an impact of factors outside of laser and scanning parameters. This is important to understand as manufacturing of parts on additive systems move towards production of parts of different sizes and build space is optimized. A quantitative understanding of the effects of secondary build variables is necessary for SLM technology.

A quantitative understanding of the effects of secondary build variables is necessary for SLM technology. The work presented here intends to address the effects of build location and build part proximity on selectively laser melted 316L stainless steel, through analysis of material structure and properties.Materials and MethodsThe samples for this research were manufactured on a Renishaw AM250, using Renishaw provided 316L stainless steel powder with a particle size distribution of 15-45 µm [13]. The Renishaw AM250 operates a 200W, 1070 nm ytterbium fiber laser with an approximate spot size of 70 µm. The laser scans the build plate and uses point exposures to selectively melt the powder layer, following the computer aided design (CAD) file. These samples were built using 50 micron layers and Renishaw’s Meander scanning parameters for 316L.

These samples were built using 50 micron layers and Renishaw’s Meander scanning parameters for 316L. The hatch fill scan parameters are: 200W laser power, a point distance of 50 microns and exposure time of 120 microseconds.Tensile samples with a radius of 4 mm were built following ASTM E8 [14] in 4 x 4 array or 5 x 5 array as seen in Figure 1. The builds were designed with different spacing between tensile samples, 10, 20, and 40 mm measured from the center of parts, to measure the effect of part proximity. Each build was produced three times for statistical significance. An additional 3 x 3 array was created with a 64 mm part spacing, wherein each part consisted of 5 tensile specimen©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA-0002839.575stacked in the z-direction as seen in Figure 1.

This conformation was only built once because of time and material limitations. All tensile samples were tested in accordance to ASTM E8 [14] under room temperature conditions. Modulus of elasticity, 0.2% offset yield strength, tensile strength, and elongation were measured. All machined gage sections followed the ASTM specification for measurement of the test diameter. Crosshead test speeds were controlled uniformly to minimize any impact of rate sensitivity and loading rates were approximately 620 MPa min-1. The test machines and extensometers meet ASTM E4 [15] and ASTM E83 [16] specifications respectively.

The test machines and extensometers meet ASTM E4 [15] and ASTM E83 [16] specifications respectively. Microstructural analysis that was completed via a scanning electron microscopy utilized electron backscatter diffraction (EBSD) images to evaluate grain size and misorientation using Orientation Imaging Microscopy™ analysis software.Figure 1: Layout of 40 mm spaced 4x4 array and layout of 3x3 array of 5 stacked tensile barsResultsThe mechanical properties of each sample were measured and analyzed by part location and spacing. The data was analyzed using two different coordinate systems in order to observe possible location based trends. Both coordinate systems were centered on the middle of the build plate. A typical orthogonal coordinate system was used, with x position increasing from left to right, and y position increasing from rear to front as shown in Figure 1.

A typical orthogonal coordinate system was used, with x position increasing from left to right, and y position increasing from rear to front as shown in Figure 1. A polar coordinate system, with radial position increasing from the center of the plate outward, and theta position following the standard unit circle was applied as well. The average yield strength, tensile strength, elastic modulus, and elongation were determined for each spacing regime by location. A 95% confidence interval was used on all of the tensile property averages, calculated using two times the standard error of the mean, assuming a normal distribution.As shown in the ultimate tensile strength results in Figure 2, the 10 and 20 mm spaced parts have an increase in strength across the x axis of the build plate. This is seen to a lesser extent in the 64 mm spaced parts.

This is seen to a lesser extent in the 64 mm spaced parts. The yield strength follows the same trend, while the elongation has the©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA-0002839.576opposite trend. This trend is not seen in the y axis position as presented in Figure 3, where there is no significant change in strength for 10 and 20 mm spaced parts. This follows for yield strength, elongation, and elastic modulus. The x and y axis position for 40 mm spacing shows a slight difference in strength dependent on the distance from the center of the build, which is supported in the radial position graph of tensile strength presented in Figure 4. This trend is not statistically present in any property other than ultimate tensile strength.

Additionally, there was no measured effect in any tensile property from z position in the stacked tensile samples as represented by tensile strength shown in Figure 5.Figure 2: Tensile strength by X axial position for each part spacing with error bars for 95% confidence intervalFigure 3: Tensile strength by Y axial position for each part spacing with error bars for 95% confidence interval©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA-0002839.577Figure 4: Tensile strength by radial position for each part spacing with error bars for 95% confidence intervalFigure 5: Tensile strength by Z position of 64 mm spaced parts with error bars for 95% confidence intervalDiscussionWhile the results show trends in the tensile performance based on location factors, it is important to note that this is limited to a small range.

The range in ultimate tensile strength was 569-631 MPa, with all the other tensile property ranges being comparably small. The trends observed are statistically significant, but limited to a small range, so in practical applications the difference may not matter.Analysis of the standard coordinate location scheme showed x position dependence, as exhibited in the 10 mm and 20 mm spaced parts in Figure 2. This was not replicated in the y position analysis, indicating the cause is a unidirectional factor. In the Renishaw AM250, Argon gas flows across the build chamber from positive to negative in the x axis. As the 10 and 20 mm spaced parts saw property changes in the x direction and not in y, it is possible that the strength change is a result of the gas flow. This makes sense as an increase in the temperature or decrease in the velocity of the gas as it flows from right to left would change convection coefficients.

This makes sense as an increase in the temperature or decrease in the velocity of the gas as it flows from right to left would change convection coefficients. As this trend was not exhibited in the parts with greater than 20 mm spacing, the thermal density of ©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA-0002839.578the build may play a role in when gas flow impacts the material properties. Parts built with tighter spacing have more concentrated thermal mass which results in slower cooling from radiation and conduction, allowing for convection to play a larger role and the gas flow to have a greater overall impact.Decreasing ultimate tensile strength with radial position was only seen in parts with 40 mm spacing, as seen in Figure 4. This trend was not seen in any other tensile property or part spacing.

This trend was not seen in any other tensile property or part spacing. It can be expected that cooling towards the edges of the build volume is faster due to faster conduction outside of the powder bed as those parts are only partially surrounded by the insulating powder bed. A faster cooling rate would most likely be associated with higher dislocation density because as a rapid change in phase is induced, the likelihood of crystallographic flaws due to volume changes is much higher. Additionally, rapid cooling will likely produce small dendritic cells as a result of constitutional undercooling, wherein the number of nucleation sites is increased. Both high dislocation density and small dendritic cells could produce higher strength.

Both high dislocation density and small dendritic cells could produce higher strength. The results in this study show lower strength in the regions of predicted faster cooling, the opposite of the expected trend in 40 mm spaced samples.Due to the lack of a radial trend seen in the 64 mm spaced parts, it appears unlikely that radial location is the single cause of trends seen in the 40 mm spaced samples. While the 64 mm spacing build required significantly more time to build, this increased time is not expected to have significant impact on the properties due to the low operating temperature of approximately 80 C. This conclusion is bolstered by the lack of any trend in tensile properties as a function of build height. Additional causes for this trend in strength could be a difference in density but the density of parts was not seen to trend radially.

Additional causes for this trend in strength could be a difference in density but the density of parts was not seen to trend radially. A second possible cause would be a difference in chemistry between parts in a build.The chemistry of the parts was analyzed using electron microscopy techniques. The chemistry of parts built with 10 mm, 20 mm and 40 mm spacing indicated no significant difference in composition as described in Table 1. Therefore the differences in tensile properties are not expected to be a result of chemical content differences.

The 64 mm spaced parts had higher manganese content, as well as lower sulfur, copper, and vanadium content, which could lead to the slightly lower strength seen in these samples.Table 1: Average and standard deviation chemical composition by weight percentCarbon Chromium Manganese Molybdenum NickelPhosphorusSulfur10 mm0.0262 (0.002)17.81 (0.091)1.20 (0.027)2.344 (0.012)12.70 (0.122)0.017 (0.002)0.010 (0.001)20 mm0.0263 (0.002)17.81 (0.087)1.193 (0.013)2.344 (0.012)12.73 (0.16)0.017 (0.001)0.010 (0.001)40 mm0.0238 (0.002)17.75 (0.117)1.188 (0.019)2.337 (0.031)12.58 (0.107)0.017 (0.001)0.010 (0.001)64 mm0.0263 (0.002)17.90 (0.061)1.274 (0.021)2.369 (0.01)12.72 (0.059)0.015 (0.001)0.007 (0.0001)©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA-0002839.579Silicon0.68 (0.018)0.693 (0.016)0.66 (0.017)0.705 (0.013)Optical microscopic images were taken and included in Figure 6.

The microstructure samples were etched with an electrolytic 70/30 nitric acid water solution. This is an image of the longitudinal surface of the part, wherein the part was sectioned to expose the XZ face. As Figure 6 shows, optical microscopy shows a dendritic structure, but one cannot clearly delineate grain boundaries from sub-grain solidification structure and therefore optical microscopy was not a suitable method for grain size characterization.

Due to this difficulty, EBSD images were taken for microstructural analysis, as large crystallographic misorientation values can be correlated to grain boundaries (Figure 7).Figure 6: Optical microscope images of 316L longitudinal surface at 2500x, scale on images, etched with an electrolytic nitric solution.Figure 7: Electron Backscatter Diffraction image of 316L longitudinal surface and inverse pole figureThe EBSD data for all parts showed no ferrite or other inclusions, indicating that all properties are a result of a single phase microstructure. A subset of data was used to evaluate the effect of part location on grain size and yield strength, wherein grain size was reported following ASTM E112 [17]. This subset contained only a few samples from each spacing, therefore statistical analysis could not be effectively performed. From the data collected, individual grains within a single sample ranged in size from less than 10 to over 100 microns.

From the data collected, individual grains within a single sample ranged in size from less than 10 to over 100 microns. Some of the observed property trends could be caused by grain size as grain boundaries block dislocation motion. The Hall-Petch relationship describes this grain boundary effect, but from the data collected, no such relationship could be determined, as seen in Figure 8. Additionally, no trend between part location and grain size could be determined. The large distribution of grain sizes could explain the difficulties in capturing representative images for analysis. The inability to identify a trend is likely from the small number of data points collected, which limits the likelihood a representative value of the mean grain size is used, and the small number of grains in each EBSD image.

©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA-0002839.580Figure 8: Grain size number versus yield strength of SLM 316L parts printed at different spacingIn addition to grain size, EBSD can provide information on the misorientation within grains which was investigated for potential relationship to part location and effect on material performance. Line scans across several grains in the EBSD images were used to measure the average sub-grain misorientation angle. This misorientation angle represents the difference in the crystallographic orientation within a grain, which can represent dendritic cell structures and the dislocation density within a grain. Again, a subset of data was analyzed and did not show a trend based on part position in build, as seen in Figure 9 for X position.

Again, a subset of data was analyzed and did not show a trend based on part position in build, as seen in Figure 9 for X position. The spread of data based on position shows no dependence in the X, Y or radial position nor part spacing, but the number of data points and lack of statistical significance means this data may not be representative of the populations mean. Additionally, there was no trend between misorientation and yield strength. The data supports relative uniformity in sub-grain misorientation, as the data is mostly limited to a tenth of degree difference.

The data supports relative uniformity in sub-grain misorientation, as the data is mostly limited to a tenth of degree difference. If this is the case, sub-grain strengthening would not explain the differences in tensile strength between spaced parts or the technique used here to represent sub- grain strengthening is not effective.Figure 9: Average sub-grain misorientation angle versus X position of SLM 316L parts printed at different spacingConclusionsA study on SLM 316L stainless steel was completed to measure the effect that part location and part spacing has on tensile properties.

The data showed there were statistically significant trends based on part location and spacing, but the range in property trends was very small, wherein ©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA-0002839.581there may be no practical, practical difference in performance. The range in ultimate tensile strength was 569-631 MPa, with all the other tensile property ranges being comparably small. It was determined that an effect of the spacing between parts in a build had an effect on tensile properties if the spacing was large enough, but below 40 mm there was no significant difference in performance. Part location was determined to have an effect related to cover gas flow for parts built at 10, and 20 mm spacing, but there was a different radial dependence in the ultimate tensile strength of 40 mm parts.

No effect was seen in the tensile properties of parts at different build heights. Preliminary analysis of chemistry, grain size and sub-grain misorientation did not provide additional support for these trends. Additional analysis is needed regarding grain size and sub- grain structure in order to provide greater statistical support for the structure and property relationships seen here. Fractography will also be done to understand the cause of failure the parts and if these are related to location.AcknowledgmentsThe Authors would like to acknowledge the assistance provided for this research by Ann Choi, David Wieliczka, and Ben Brown.All data prepared, analyzed and presented has been developed in a specific context of work and was prepared for internal evaluation and use pursuant to that work authorized under the referenced contract.

Reference herein to any specific commercial product, process or service by trade name, trademark, manufacturer, or otherwise, does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government, any agency thereof or Honeywell Federal Manufacturing & Technologies, LLC.This presentation has been authored by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA-0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.Conflict of Interest: The authors declare that they have no conflict of interest.Bibliography[1] N. Hrabe and T. Quinn.

[8] P.Y. Peretyagin, et al. “Track Geometry in Selective Laser Melting,” Russian Engineering Research, 35 (2015): 473-476. [9] B. Song et al, “Differences in microstructure and properties between selective laser melting and traditional manufacturing for fabrication of metal parts: a review,” Mechanical Engineering, 10 (2015): 111-125. [10] A.S. Wu et al. “An Experimental Investigation into Additive Manufacturing Induced Residual Stresses in 316L Stainless Steel,” Metallurgical and Materials Transactions A, 45 A (2014): 6260-6270. [11] H . K. Rafi, N. V. Karthik et al. “Microstructures and Mechanical Properties of Ti6Al4V Parts Fabricated by Selective Laser Melting and Electron Beam Melting,” Journal of Materials Engineering and Performance, 22 (2013): 3872-3884. [12] L. Ladani. “Local and Global Mechanical Behavior and Microstructure of Ti6Al4V Parts Built Using Electron Beam Melting Technology,” Metallurgical and Materials Transactions A, 46(9) (2015): 3835-3841.

Ladani. “Local and Global Mechanical Behavior and Microstructure of Ti6Al4V Parts Built Using Electron Beam Melting Technology,” Metallurgical and Materials Transactions A, 46(9) (2015): 3835-3841. [13] Renishaw, 2015, “SS 316L-0407 powder for additive manufacturing,” Renishaw plc, Stone, Staffordshire, UK, 2015. H-5800-3001-01-D, www.renishaw.com/additive. [14] ASTM Standard E8, 2015, “15a Standard Test Methods for Tension Testing of Metallic Materials,” ASTM International, West Conshohocken, PA, 2015, DOI: 10.1520/E0008-15A, www.astm.org. [15] ASTM Standard E4, 2014 “Standard Practices for Force Verification of Testing Machines,” ASTM International, West Conshohocken, PA, 2014, DOI: 10.1520/E0004-14, www.astm.org. [16] ASTM Standard E83, 2010, “Standard Practice for Verfitication and Clasfficiation of Extensometer Systems,” ASTM International, West Conshohocken, PA, 2010, DOI: 10.1520/E0083-10A, www.astm.org.

[16] ASTM Standard E83, 2010, “Standard Practice for Verfitication and Clasfficiation of Extensometer Systems,” ASTM International, West Conshohocken, PA, 2010, DOI: 10.1520/E0083-10A, www.astm.org. [17] ASTM Standard E112, 2014, “Standards Test Methods for Determining Average Grain Size,” ASTM International, West Conshohocken, PA, 2014, DOI: 10.1520/E0112-13., www.astm.org.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA-0002839.583.

Austin, TX 78735, United States3Henry Samueli School of Engineering and Applied Science, University of California, Los Angeles, CA 90095, USAAbstractMetal nanoparticles have been extensively investigated for the applications in micro-nano electronics,plasmonics, and chip packaging due to their enhanced thermal, mechanical and electrical properties at the nanoscale.For instance, surface-to-volume ratio and particle distribution significantly affect the sintering and melting behavior ofnanoparticles relative to bulk properties. Hence, understanding the characteristic behavior of these materials is vital forthe use of these nanoparticles in existing and new applications.

Hence, understanding the characteristic behavior of these materials is vital forthe use of these nanoparticles in existing and new applications. It has been shown that at the nanoscale van der Waalsforces are the dominant force between particles, which results in particle agglomeration and creates particle clustering.In this paper, we show how this nanoparticle clustering effects the optoelectronic property of the domain by applyingthe finite difference frequency domain method.

This is the first step towards understanding the modeling of the heattransfer that characterizes the laser particle interaction for additive manufacturing using nanoparticles.Keywords: Nanotechnology, Surface Plasmon Polaritons, Near-Field Energy Enhancement1.IntroductionThe industrial demand for smaller devices have resulted in the development of nanomanufacturing.There has been suggested some methods such as lithography; however, most of these methods are restrictedby the diffraction limit of the light and their inability of large area manufacturing. Hence, the development ofnew methods for metals and ceramics are vital to fabricate additively-manufactured parts with sub-micronscale resolutions which makes the harnessing the heat transfer at the nanoscale become crucial to overcomingthe diffraction limit by the creation of near-field heat transfer around the structures [1,2].

There has been sometechniques based on near-field optics have been applied to confining the light at the nanoscale. Nanomaterialssuch as metal nanoparticles have attracted interest as their optical, thermal, and electromechanical propertiesof nanomaterials are quite different from the properties of the bulk material, and they can be tuned easily bycontrolling the shape and size of the nanostructure. The noble metals have high electron mobility and areconsidered good electrical conductors, making them attractive candidates for plasmonic applications. Theysupport the localized surface plasmons (LSPs) and surface plasmon polaritons (SPPs) due to their ability to1341enhance the electromagnetic fields and light confinement between metal and dielectric materials. Theinteraction or coupling of the plasmons on the nanoparticles provide practical strategies to the design ofnanostructures.

Theinteraction or coupling of the plasmons on the nanoparticles provide practical strategies to the design ofnanostructures. Also, the surrounding dielectric medium changes the plasmonic properties of the nanoparticleswhich has been big interest in the development of plasmon resonants [3,4]. The resonances have resulted insome exciting applications such as Surface Enhanced Raman Scattering (SERS) and nanoscale sensing, whichis manifested by the enhanced scattering and absorption cross sections of nanoparticles. However,understanding of plasmonic nanoparticle clusters is a significant challenge. Recently, metal nanoaparticleshave been used in the additive manufacturing of electronics, aerospace and medical device parts throughmicroscale selective laser sintering (μ-SLS) where nanoparticles replace the traditional powders used instandard SLS processes [5, 6]. Hence, many efforts have been devoted to the fabrication of metallicnanostructures by different methods [7].

Hence, many efforts have been devoted to the fabrication of metallicnanostructures by different methods [7]. Different shapes and sizes have been synthesized and plasmonicproperties have been studies previously [8]. The plasmonic resonance behavior of some anisotropicnanostructures show different behavior than the spherical metal nanoparticles. For example, the plasmonicresonance peaks of anisotropic nanostructures move towards longer wavelengths, and some anisotropicallyetched nanoparticles show broad bandwidths [9].The cohesive interactions between the nanoparticles, dominate interactions between the nanoparticleat nanoscale can result in agglomeration [10, 11]. These agglomeration effects can significantly reduce thepacking density of the nanoparticles and show different field enhancement and plasmonic effects comparedto the individual metallic nanoparticles.

Hence, optoelectronic properties of such agglomeration effects onplasmonic behavior need to be understood well so to be tuned accurately by changing the gap distances andthe size of the adjacent nanoparticles [12]. In this paper, we investigate the plasmonic properties of sphericalcopper nanoparticles with agglomeration effect using finite different frequency domain method. We displayedthe near-field intensity enhancement and the movement of the LSPR peaks of the Cu nanoparticles.2. Modeling ApproachFinite element method shows higher accuracy for modeling the irregular shapes compared to othermethods and enables to simulate finite elements within large domain. Khoury et al. have revealed the validityand the accuracy for plasmonic-nanostructure modeling by finite element model [13].

Khoury et al. have revealed the validityand the accuracy for plasmonic-nanostructure modeling by finite element model [13]. In this study, finitedifference frequency domain based on finite element method (FEM) has been used to understand near-fieldintensity distribution of metallic nanostructures by solving Maxwell’s equations. Moreover, the Cunanoparticles were surrounded with air which the refractive index is taken as n=1 and complex permittivity2342of the metallic nanoparticle is used from experimental data from [14].

Also, non-uniform meshes areimplemented effectively and produce convergence as well as saving computational time [15].As the nanoparticles exhibit different plasmonic effects at different size and gap distances,configuration of the particles need to be understood well to be able to tune the optical properties of plasmonics.As the metallic nanoparticles agglomerate at the nanoscale, we used two different configurations in whichthere are three particles initially are agglomerated in both cases. In the first case, we used linear configurationand in the second case, we analyzed the triangular configuration. The configurations are illustrated in Fig. 1.Figure 1. Linear (Case 1) and Triangular (Case 2) ConfigurationsIn this paper, we have analyzed two cases in which the nanoparticles are placed in linear and triangularconfigurations.

1.Figure 1. Linear (Case 1) and Triangular (Case 2) ConfigurationsIn this paper, we have analyzed two cases in which the nanoparticles are placed in linear and triangularconfigurations. We have used copper nanoparticles with 100 nm diameter on a 350 nm thick, 750 nm x 750nm wide glass substrate.To be able see the agglomeration effect, we have varied the distance between theparticles (d) from 0 to 50 nm. By investigating TE polarized light, scattering and absorption cross sections arecalculated. Scattering cross section is found by calculating the surface integration of the scattered Poyntingvector and the absoprtion crosssection is analyzed by the volume integration of the absorbed energy withinthe nanoparticles. 532 nm wavelength, TE polarized light is irradiated for each cases.3. Results and DiscussionFig. 2 shows the linear (Case 1) configuration electric field enhancement for d=0 at the center of thenanoparticles.

Results and DiscussionFig. 2 shows the linear (Case 1) configuration electric field enhancement for d=0 at the center of thenanoparticles. It indicates that surface plasmons are generated along the polarization direction of thenanoparticles with up to six times that of incident electric field and up to seventeen times at the interfacebetween the particles. This shows that strong localized field is generated below the diffraction limit.3343Figure 2. Top View of Electric Field Enhancement of Case 1for d=0 and λ=532 nmFigure 3. Side View of Fig 2.Fig 4. illustrates the linear (Case 1) configuration in which the distance between the particles are variedfrom 10 nm to 50 nm. It can be seen that the plasmonic effects are observed only on the polarization directionand up to 6 times of that incident electric field.Fig. 5 depicts the triangular (Case2) configuration when d=0.

5 depicts the triangular (Case2) configuration when d=0. It is noticed that the electric field isenhanced up to 120 times between the adjacent nanoparticles, which is around 7 times higher than the linearconfiguration when d=0. We also observed that the maximum enhancement occur between the adjacentnanoparticles and due to the polarization direction we don’t see the enhancement between the two adjacentnanoparticles which is located at the bottom of the figure in Fig. 5. Also, we see the plasmonic effect is on thedirection of the polarization on the particles.4344Figure 4. Electric Field Enhancement of Case 1 with varying dFigure 5. Electric Field Enhancement of Case 2 for d=05345Figure 6. Electric Field Enhancement of Case 2 with varying dFig. 6 illustrates the electric field enhancement with varying d from 10 nm to 50 nm. It can be seenthat the localized field enhancement is up to 12 times when d is 10 nm.

6 illustrates the electric field enhancement with varying d from 10 nm to 50 nm. It can be seenthat the localized field enhancement is up to 12 times when d is 10 nm. Moreover, the surface plasmonpolariton is much stronger within the particles and show a narrow excitation between the nanoparticles. Thislocalized enhacement is decreased to between 6 to 7 times when d is between 20 nm to 50 nm. Anotherinteresting result is that strength of the near-field effect is shifting from the between of the particles to thesurface of the particles when d is between 20 to 50 nm and at 50 nm the particles start to behave almostindividually. Also, the plasmonic effect can be seen along the direction of the polarization only (TE polarized)when d is at 50 nm. This indicates that the light/nanostructure interaction is affected significantly for differentdistance values between the particles as well as different configurations.

This indicates that the light/nanostructure interaction is affected significantly for differentdistance values between the particles as well as different configurations. It can be observed that the enhancedfield is always much higher for triangular configuration than the linear case when the particles form clusterwhich means when d is 0. This enhanced field drop significantly when the particles start to have some spacingbetween each other. This gives an insight that the energy transfer between the particles are affectedsignificantly for different configurations when they are at cluster form or d=0.

This gives an insight that the energy transfer between the particles are affectedsignificantly for different configurations when they are at cluster form or d=0. Another interesting observationis that when the particle’s distance, d , is 50 nm, the field enhancement is almost the same for linear andtriangular configuration which implies that particles start to lose their interaction between each other at thenear-field regime and behave similarly for these configurations.6346As the light is absorbed and scattered by the metallic nanoparticle at 532 nm wavelength, absorptionand scattering crosssection is important to understand for two cases that we looked at and mentioned in theprevious section. It can be seen from Fig. 7 that absorption increases with particle spacing for linear case;however, absorption starts to gradually decrease with spacing for triangular case.

It can be seen from Fig. 7 that absorption increases with particle spacing for linear case;however, absorption starts to gradually decrease with spacing for triangular case. This can be explained as theinterference of propagation around the triangle starts to be reflected at the triangle corners. We can also seethat at large spacing; the particles act more as individual particles and the results converge.Figure 7. Absorption and Scattering Crosssection (m2) versus particle spacing (d)Also, we can see that scattering increases with particle spacing for both cases. This increase is linear for case 1 whichis linear configuration. For case 2 which is the triangular configuration, interference between background and a resonantscattering results in asymmetric line shape or Fano type resonance. This phenomena is caused by the interferencebetween two scattering amplitudes within the background and the resonant process.

This phenomena is caused by the interferencebetween two scattering amplitudes within the background and the resonant process. Indeed, this effect is observed whend is between 0-10 nm in which the near-field energy is localized and create an almost coherent excitation.Fig. 8 shows the extinction crosssection which is the summation of absorption and scattering crosssection. Itcan be seen that scattering is more dominant over absorption at 100 nm of copper nanoparticle when interacts with 532nm of wavelength and TE polarized light and thus the extinction crosssection is characterized by mostly scatteringeffect. We can also see that the two cases converge at the large particle spacing.As it can be seen that the absorption crosssection is always higher for triangular case than the linearconfiguration. Moreover, scattering crosssection is higher for linear case.

Moreover, scattering crosssection is higher for linear case. This also shows that 532 nm of wavelengthand TE polarized light does not penetrate as far into the linear configuration compared to the triangular configuration.This implies that how far the energy is penetrated through the nanoparticles which affects the depth of the sintering in7347the powder bed. Similarly, the closer the particles are together, the less light propagates into the powder bed which alsoaffects the sintering depth.Figure 8. Extinction Crosssection (m2) versus particle spacing (d)4. ConclusionWe have shown that solutions of Maxwell’s equation in frequency domain can depict the ability ofnear-field enhancement which results in surface plasmon generation below the diffraction limit. We examinedthe spacing effect between the nanoparticles with linear and triangular configurations for this behavior andobserved the field enhancement to be highly dependent on the nanoparticle configuration.

It is observed thatnanoparticles when they are in a cluster group; in other words, when they are all touching each other, themaximum enhancement is achieved compared to the cases where there is a gap between the particles. Thethermal radiation properties are nonlinearly characterized from 0-10 nm regime due to the strong and non-local near-field interaction between the particles with the linear and triangular configuration cases analyzedin this study. That shows that the light interaction with nanoparticles can be tuned by analyzing differentconfigurations and the optimum configuration could be satisfied to achieve desired thermal radiationparameters to understand the sub-wavelength energy transport.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEFFECT OF OPTIMIZING PARTICLE SIZE IN LASER METAL DEPOSITION WITH BLOWNPRE-MIXED POWDERSaW. Liaa , J. W. Zhanga , X. C. Zhanga , S. Karnatia , F. Liou.Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409, United StatesAbstractFunctionally Graded Material (FGM) is often fabricated by Laser metal deposition with pre-mixed multiple powders (PMM-powder). Since the supplied PMM-powder directly affects FGM’s composition, investigation on PMM-powder’s property is greatly needed. This paper employed experimental method to observe an important problem: PMM-powder separation in fabricating FGM. A novel particle size optimization method was introduced as solution to eliminate the powder separation.

A novel particle size optimization method was introduced as solution to eliminate the powder separation. Pre-mixed pure Cu and 4047 Al powders were used to do two experiments. The first experiment result disclosed the existence of powder separation. By optimizing the particle size, the PMM-powder separation was effectively solved in the second experiment result.Keywords: Functionally Gradient Material; Pre-Mixed Multi-Powder; Laser metal deposition; powder separation; composition deviation. 1. IntroductionFunctionally Gradient Material (FGM) is arousing more and more attention in area of advanced material processing. Gradual variation in composition and structure over volume in FGM [1], result in corresponding gradual changes in the properties of the material to meet the usage requirements. FGM can be designed for specific functions and applications, and can combine two or more different materials into one structure and perform multiple functions [2, 3].

FGM can be designed for specific functions and applications, and can combine two or more different materials into one structure and perform multiple functions [2, 3]. To fabricate the FGM structure, laser metal deposition is an effective process technique, in which gradually varying in composition over volume is achieved by Pre-Mixed Multi-Powder (PMM-powder) [4, 5]. Fig.1 is schematic illustration of fabricating FGM by laser metal deposition.234Fig.1. Schematic illustration of fabricating FGM by Laser metal deposition.A necessary factor in FGM fabrication is the multi-composition powder, which is often pre-mixed of two kinds of alloy or metal elemental powders according to the composition requirement of FGM structure. In the process of Laser metal deposition, the PMM-powder are input into powder feeder. Inert argon gas flow drives the powders to move along the powder feeder pipe.

In the process of Laser metal deposition, the PMM-powder are input into powder feeder. Inert argon gas flow drives the powders to move along the powder feeder pipe. Since the PMM-powders have different particle densities and sizes, under the same argon gas flow, light and small particles could move faster while heavy and big particles could lag behind. This kind of particle flow dynamics may cause powder separation and composition deviation in PMM-powder. In FGM structure, the material composition ratio in any different area is required strictly (Fig.1). Non-uniform PMM-powders will ruin the gradually varying material composition ratio, and further discount the wanted material performance. Therefore, to address this issue, investigation on PMM-powder during laser metal deposition is greatly needed. So far, there is no special study focusing on pre-mixed multiple powders’ dynamic behavior in fabricating FGM by laser metal deposition.

So far, there is no special study focusing on pre-mixed multiple powders’ dynamic behavior in fabricating FGM by laser metal deposition. Some previous research results about the powder flow were just for identical material through both numerical and experimental methods [6-10].In this paper, experimental method was employed to observe powder separation in fabricating FGM with blown PMM-powders. A novel particle size optimization method was introduced as solution to eliminate the powder separation. Pre-mixed pure Cu and 4047 Al powders were used to do two experiments. The first experiment result disclosed the existence of powder separation. By optimizing the particle size in two types of powder, the powder separation was effectively solved in the second experiment result. 2. Experiment procedure2.1. Experiment set-up235The experiment set-up is schematically shown in Fig.2.

2. Experiment procedure2.1. Experiment set-up235The experiment set-up is schematically shown in Fig.2. A commercial powder feeder (Bay State Surface Technologies, Inc, Model-1200) was used to supply PMM-powder. Pure Cu powder and 4047 Al powder (Fig.2) were used in the experiments. The particles size distributions were displayed by the sieve analysis in Table.1 and Table.2. A plastic pipe, with 1.5 m in length and 5 mm in inner diameter, delivers the PMM-powder. Two linear motors were employed to generate moving path. A piece of 6061 Aluminum alloy plate, paved with sticky epoxy resin, was fixed on the linear motor to follow the generated path. The idea of this experiment is to use the sticky epoxy resin layer to collect the pre-mixed 4047 Al and Cu powders spraying out from the nozzle. The optical microscopy was used to observe the PMM-powder adhered on epoxy resin layer. The time of spraying powder was 25 s.

The optical microscopy was used to observe the PMM-powder adhered on epoxy resin layer. The time of spraying powder was 25 s. During this time range, powder feeder nozzle was moving above the epoxy resin layer following the specified path “M”. After the epoxy resin was solidified, the collected PMM-powder was contained in it to form the particle pattern, then for the further observation.Table.1. Sieve analysis of pure Cu powderSieve type70 mesh100 mesh120 mesh140 mesh200 mesh325 meshSize (μm)>212150-212125-150106-12575-10645-75Percentage (%)0.01.32.43.747.445.2Table.2. Sieve analysis of 4047 Al powderSieve type70 mesh100 mesh120 mesh140 mesh200 mesh325 meshSize (μm)>212150-212125-150106-12575-10645-75Percentage (%)1.12.85.420.342.527.92362.2. Experiment design ideaFig.2. Experiment set-up and results.This research included two experiments.

Experiment design ideaFig.2. Experiment set-up and results.This research included two experiments. The experiment design idea was based on the thinking of particle acceleration equation in Eqn.2, which is a classic first order differential equation with variable of particle velocity 𝑢𝑢𝑝𝑝. The second term on the right side can be simplified as gravitational acceleration 𝑔𝑔 because particle density 𝜌𝜌𝑝𝑝 is much greater than argon gas density 𝜌𝜌. The first term on the right side contains several coefficients including argon gas flow velocity 𝑢𝑢, drag coefficient 𝐶𝐶𝐷𝐷, argon gas density 𝜌𝜌, Reynolds number 𝑅𝑅𝑅𝑅, and the product of particle diameter 𝑑𝑑𝑝𝑝 and density 𝜌𝜌𝑝𝑝.

Much closer the 𝑑𝑑𝑝𝑝_𝐴𝐴𝐴𝐴/𝑑𝑑𝑝𝑝_𝐶𝐶𝐶𝐶 is to the ideal value, more effectively the PMM-powder separation could be eliminated. Otherwise, if the 𝑑𝑑𝑝𝑝_𝐴𝐴𝐴𝐴/𝑑𝑑𝑝𝑝_𝐶𝐶𝐶𝐶 much greater or less than the ideal value, the PMM-powder will separate when flowing in powder feeder pipe since the mixed particles have different acceleration equations.Based on above analysis, two experiments were designed. These two experiments were performed with same operating parameters, but different particle diameter. In experiment-1, the 𝑑𝑑𝑝𝑝_𝐴𝐴𝐴𝐴/𝑑𝑑𝑝𝑝_𝐶𝐶𝐶𝐶 was much less than the ideal value. But in experiment-2, the 𝑑𝑑𝑝𝑝_𝐴𝐴𝐴𝐴/𝑑𝑑𝑝𝑝_𝐶𝐶𝐶𝐶 is much close to the ideal value. The detailed operating parameters in the experiments were shown in Table.3.Table.3.

But in experiment-2, the 𝑑𝑑𝑝𝑝_𝐴𝐴𝐴𝐴/𝑑𝑑𝑝𝑝_𝐶𝐶𝐶𝐶 is much close to the ideal value. The detailed operating parameters in the experiments were shown in Table.3.Table.3. Sieve analysis of pure Cu powderExperiemnt-1Experiment-2(optimized)Schematic particles (Gray-Al; Red-Cu)4047 Al particle diameter Cu particle diameter 𝑑𝑑𝑝𝑝_𝐴𝐴𝐴𝐴/𝑑𝑑𝑝𝑝_𝐶𝐶𝐶𝐶45-75 μm 75-106 μm 0.667 23775-106 μm 45-75 μm 1.5Vol.% ratio Argon gas flow rate Al plate moving speed Moving time50%:50% 6 m/s 1m/min 25 s50%:50% 6 m/s 1m/min 25 s3. Results and Discussion3.1. Analysis methods for the experiment resultsFig. 2 shows the two solidified epoxy resin layers containing PMM-powder patterns from the two experiments. It was found that the first side in “M” shape was lost in the experiments, because when the nozzle moving along the first section of “M”, the powder flow was right in the pipe.

It was found that the first side in “M” shape was lost in the experiments, because when the nozzle moving along the first section of “M”, the powder flow was right in the pipe. 87 observed zones with the size of 4mm by 3mm distributed with uniform-interval along the central line of the powder path for microscopic particle patterns observation. Then the distribution of different particles in the pattern was quantified. Due to the color difference of PMM-powder, the image processing software Image-J was used to mark 4047 Al particles and Cu particles in each observed zone, then count the particle numbers of two powders. Since particle sizes were known by the sieve analysis, the volume percentage of two powders in each observed zone can be calculated by Eqn.5 and Eqn.6.4 3 𝑉𝑉𝑉𝑉𝑉𝑉𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝13 𝜋𝜋𝑟𝑟𝑚𝑚𝑝𝑝𝑝𝑝𝑚𝑚𝑉𝑉𝑉𝑉𝑉𝑉𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝 = 𝑁𝑁𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝐴𝐴𝑝𝑝 ×(5)𝑉𝑉𝑉𝑉𝑉𝑉. %𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝1 = 100 ×𝑉𝑉𝑉𝑉𝑉𝑉𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝1 + 𝑉𝑉𝑉𝑉𝑉𝑉𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝2 (6)3.2. Results of experimentsFig.3.

%𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝1 = 100 ×𝑉𝑉𝑉𝑉𝑉𝑉𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝1 + 𝑉𝑉𝑉𝑉𝑉𝑉𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝𝑝2 (6)3.2. Results of experimentsFig.3. Observation of the experiment results.238The quantification of PMM-powder patterns from experiment-1 was shown in Fig.4a. Blue color recorded the volume percentage of 4047 Al, while red color recorded the volume percentage of Cu. The time in the X-axis was used to indicate the moments when the particles in the relative observation zone reached the epoxy resin. From the PMM-powder history, a clear tendency can be found. In the beginning stage (approximately 0s ~ 4s), no particle sprayed out from the nozzle, so the volume percentage of two powders was zero. Starting from the moment of 4s, a flow of PMM-powder exited from the nozzle, in which 4047 Al was the primary particles, whose volume percentage reached to 93.25%. Some rare red Cu particles were found, whose volume percentage was 6.75%.

Some rare red Cu particles were found, whose volume percentage was 6.75%. With time increasing, more particles sprayed out from powder feeder nozzle, the primary particles are still 4047 Al, but whose volume percentage was decreasing gradually. In stark contrast, the volume percentage of Cu was gradually increasing. At the moment of 21s, Cu and 4047 Al had approximately same volume percentage. Then from the moment of 21s to the end, Cu powder volume percentage surpassed the 4047 Al powder volume percentage. Since both powders started to transit simultaneously from powder feeder pipe inlet, and 4047 Al particle had bigger acceleration than Cu particle according to the Eqn. 2, 4047 Al moved faster and exited from nozzle earlier than Cu particle. This is why Al powder volume percentage was greater than Cu powder.

2, 4047 Al moved faster and exited from nozzle earlier than Cu particle. This is why Al powder volume percentage was greater than Cu powder. The powder volume percentage history in Fig.4a implied the PMM-powder separation since the two powders’ volume percentage were not uniform and kept changing, although they were designed as 50%:50%. The PMM-powder separation in their flow movement ruined the powder design and resulted in the severe composition deviation between designed FGM and real deposited FGM.The quantification of PMM-powder composition patterns in experiment-2 was shown in Fig.4b. By observing the optimized PMM-powder’s composition distribution patterned in the epoxy resin coating, the PMM-powder flow had basically constant volume percentages, which were close to 50%:50%. Comparing with the results in experiment-1, the powder separation was effectively eliminated, so that the composition deviation was reduced.(a)239(b)Fig.4.

Comparing with the results in experiment-1, the powder separation was effectively eliminated, so that the composition deviation was reduced.(a)239(b)Fig.4. PMM-powder volume percentage history in (a)experiment-1, and (b)experiment-2.4. ConclusionSome conclusions in this paper were summarized as follows. In the PMM-powders, which have two types of powder with different densities, if the particle diameter ratio is equal to or close to an ideal value, all the particles will have the same acceleration when moving in powder feeder pipe. The ideal value is the square root of powder density ratio. Therefore, there will be very less powder separation. The designed composition in PMM-powder can be retained.Otherwise, if the particle diameter ratio is much greater or less than the ideal value, powder separation will happen in powder feeder pipe due to the different particle accelerations. It will cause the deviation between designed composition and real composition.

98-104. [6] Pan, H., Sparks, T., Thakar, Y. D., and Liou, F., 2006, "The investigation of gravity-driven metal powder flow in coaxial nozzle for laser-aided direct metal deposition process," Journal of manufacturing science and engineering, 128(2), pp. 541-553. [7] Pan, H., and Liou, F., 2005, "Numerical simulation of metallic powder flow in a coaxial nozzle for the laser aided deposition process," Journal of Materials Processing Technology, 168(2), pp. 230-244. [8] Tan, H., Zhang, F., Wen, R., Chen, J., and Huang, W., 2012, "Experiment study of powder flow feed behavior of laser solid forming," Optics and Lasers in Engineering, 50(3), pp. 391-398. [9] Pinkerton, A. J., and Li, L., "A verified model of the behaviour of the axial powder stream concentration from a coaxial laser cladding nozzle," Proc. Proceedings of ICALEO.

391-398. [9] Pinkerton, A. J., and Li, L., "A verified model of the behaviour of the axial powder stream concentration from a coaxial laser cladding nozzle," Proc. Proceedings of ICALEO. [10] Wen, S., Shin, Y., Murthy, J., and Sojka, P., 2009, "Modeling of coaxial powder flow for the laser direct deposition process," International Journal of Heat and Mass Transfer, 52(25), pp. 5867-5877.241.

With the aim to improve the flowability of grinded powders and, thus, the SLS processability, a thermal process for particle shape modification was developed. This has been applied successfully for a broad variety of thermoplasts. In this contribution, a modified design of a rounding reactor with direct heating and optimized geometry is presented. The performance of the new design is assessed by the rounding of commercially available polypropylene powders. The influence of the rounding process on the packing density, flowability and particle size and shape distributions of the powders is investigated. Test specimens produced by SLS were tested to assess the effect of the rounding process on the mechanical properties of the produced parts.IntroductionSelective laser sintering (SLS) of polymers is one of the additive manufacturing (AM) techniques, which allows the production of 3D functional parts with complex geometries without the need for tooling and assembly [1].

During the manufacturing process, consecutive layers of powder are applied with a blade or a roller and then sintered with a laser, before the next layer of powder is applied [2–4]. The mechanical properties and surface quality of the produced devices depend to a great extent on the intrinsic material (e.g. melting behavior, crystallization) and powder bulk (e.g. particle size and shape, flowability) properties [5–7]. Desirable powders are of narrow particle size distribution with volume-averaged mean particles sizes between 15 and 100 μm made of spherical or potato-shaped particles [8]. The choice of commercially available SLS materials is rather limited. Currently, the most widely used polymeric material for SLS is polyamide 12 (PA12) with a market share of around 90% [9]. Besides PA12, other types of polyamides (PA11 and PA6), polystyrene, polypropylene and polyether ether ketone are available in minor extent [9].

Besides PA12, other types of polyamides (PA11 and PA6), polystyrene, polypropylene and polyether ether ketone are available in minor extent [9]. One popular process for the production of polymer powders for SLS is comminution [10,11]. Nevertheless, comminuted particles normally have an irregular shape with different aspects ratios, which results in powders with low flowability and packing density [11–13]. To improve the flowability of these powders, an additional step of shape modification is required. With this aim, a thermal process for shape modification was developed [10]. This process consists of the rounding of irregular polymer particles in a heated downer reactor. This process has been applied successfully for a broad variety of polymers [10,14,15], nevertheless, low yields due to particle build-up at the reactor walls [16], and agglomeration are the main drawbacks of the process. In this study, a modified rounding reactor concept is presented.

In this study, a modified rounding reactor concept is presented. The efficiency of the modified concept in terms of yield is analyzed by considering the rounding of commercial790polypropylene powders. Furthermore, the effect of the rounding process on the bulk properties and processability of polymer powders is investigated. With this aim, the untreated and rounded powders are characterized according to shape and size distribution, as well as packing density and flowability. Test specimens produced by SLS of the rounded and modified powders were tested to assess the effect of the rounding process on the processability of the powders and the mechanical properties of the produced parts.MethodsRounding reactorA batch of 600 g of polypropylene (PP) powder (Coathylene® PD0580, Axalta polymer powders, Switzerland) was rounded in the downer reactor.

The downer setup is composed of a stainless steel pipe (inner diameter =100 mm, length=600 mm), an aerosol generator unit, a cooling section, and a separation system to recover the rounded particles. The powder feed is dispersed in nitrogen (5.0, Air Liquide Deutschland GmbH, Germany) by means of a flat-tray feeder (ZD 22 FB-C-2M, Three-tec GmbH, Germany) equipped with a self- made powder venturi injector at a pressure of 6 bars and a flow rate of 6 Nm3/h. In the current study, the powder feed rate was 100 g/h. The aerosol flow is heated to 120°C by using a heating band (HBS, Horst GmbH, Germany) wrapped around the aerosol pipe, which is centrally fed into the head of the reactor. A nitrogen sheath flow is heated to 210°C employing a gas heater (GA00565, Horst GmbH, Germany). The function of the sheath flow is to provide the energy necessary to melt the particles and to reduce the interactions of the particles and the wall [16].

The function of the sheath flow is to provide the energy necessary to melt the particles and to reduce the interactions of the particles and the wall [16]. The volumetric flow of the sheath gas was set at 7 Nm3/h by a mass flow controller (EL-Flow, Wagner Mess-und Regeltechnik, Germany). The sheath gas is homogeneously distributed over the downer cross-section surrounding the aerosol gas by means of a sintered metal plate (SIKA-R20, GKN Sinter Metals GmbH, Germany). The molten particles acquire progressively a spherical shape driven by the minimization of surface energy (c.f. minimization of free surface) of the melt droplet. At the end of the reactor, the gas and particles are cooled by quenching with air at room temperature. Finally, the rounded solid particles are separated from the gas flow by means of a deflecting separator and cyclone.

Finally, the rounded solid particles are separated from the gas flow by means of a deflecting separator and cyclone. After the experiment, the mass of the product obtained was measured to determine the yield of the process.Scanning electron microscopyParticle shape and surface morphology were characterized by scanning electron microscopy (SEM). A Gemini Ultra 55 (Zeiss) was used to take images of the coated particles using an SE2 detector at an acceleration voltage of 1kV.Laser diffraction particle sizingParticle size distribution (PSD) was determined by laser diffraction (Mastersizer 2000, Malvern Panalytical GmbH, Germany) according to ISO13320-1, using the dry dispersion unit Scirocco 2000 at 50% of the vibration amplitude of the feeding tray.Light microscopyLight microscopy was also used to determine the particle size and shape distribution of the feed and rounded powders.

With this aim, a few milligrams of powder were applied on a microscope slide and were spread by friction with the help of another microscope slide, in order to reduce particle agglomeration. Images were taken by using an Axio Imager M1m light microscope (Carl Zeiss Microscopy GmbH, Germany) in transmitted light mode with a 10x objective. The motorized microscope stage allows to scan the complete slide generating an array of images, which were digitally analyzed in ImageJ software in order to determine the particle size and shape factor distributions of the powders. For the evaluation, at least 8000 particles were analyzed.

For the evaluation, at least 8000 particles were analyzed. The shape factor used in the image analysis was the circularity defined as [19]791(cid:1829)(cid:1861)(cid:1870)(cid:1855)(cid:1873)(cid:1864)(cid:1853)(cid:1870)(cid:1861)(cid:1872)(cid:1877) = 4 (cid:2024) (cid:1499)(cid:1827)(cid:1870)(cid:1857)(cid:1853) (cid:1842)(cid:1857)(cid:1870)(cid:1861)(cid:1865)(cid:1857)(cid:1872)(cid:1857)(cid:1870)(cid:2870), where a value of 1 indicates a perfect circle and a value nearing 0 represents a very elongated particle.FlowabilityA ring shear tester (RST-01.01, Dr. Dietmar Schulze Schüttgutmesstechnik, Germany) was used to measure the flow behavior of the plasma-treated powder. The powders are filled in the tester and the measuring device was prepared according to procedures indicated in ASTM D6773 – 02 [17]. The powders were subjected to consolidation stresses of 1450 Pa, 2500Pa and 5800 Pa.

The flow function provides a measure of the flowability of the powder at a defined consolidation stress. In general, higher values of ffC represent good flowability, while small values characterize cohesive materials. A value in the range of 1-2 corresponds to very cohesive, while a value larger than 4 to an easily or free flowing powder. The reported values of ffc correspond to the mean value of three separated measurements [18].Packing densityThe packing density of the samples was determined by filling the powders into an aluminum cylinder with a previously calibrated volume of 243.46 mL. The cylinder was filled with the powders using a spoon until the powder is over the top of the cylinder wall in a way, that (consolidation) forces were omitted as far as possible. The excess of material was removed by scrapping the cylinder with a blade. The weight of powder filled in the cylinder was determined and the (loose) packing density was calculated.

The excess of material was removed by scrapping the cylinder with a blade. The weight of powder filled in the cylinder was determined and the (loose) packing density was calculated. Measurements were repeated three times for each sample. Mean values and standard deviations are reported.Laser sinteringTo investigate the influence of the rounding on the SLS processability and mechanical properties of the produced devices, test specimens from the untreated and rounded powders were produced in a SnowWhite desktop SLS machine (Sharebot, Italy). The small building chamber (100x100x100 mm3) allows for production of tensile strength specimens of type 5A according to DIN EN ISO 527-2 [20]. The geometry of the test specimens was sliced in 10 layers of 200 μm by using the freeware software Slic3r 69. The temperature in the machine was controlled by using the “Environment” mode, which controls the heaters by the environment temperature in the machine.

The temperature in the machine was controlled by using the “Environment” mode, which controls the heaters by the environment temperature in the machine. The environment temperature was set to 126°C, which equals a powder bed temperature of 148°C. The laser power and scanning speed were set to 2.8 W and 1100 mm/s. The powder reservoir was elevated 290 μm after each layer in order to obtain homogenous and closed layers of powder. Six test specimens were produced from each powder sample.Characterization of the dimension stability and mechanical properties of test specimensAll tensile tests of SLS produced specimen were performed on a Z050 (Zwick) tensile testing machine according to DIN EN ISO 527-1 [21]. The test speed was 20 mm/min, the clamping length 50 mm and the preload 0.1 N. The Young’s modulus was determined at a speed of 0.5 mm/min with an elongation between 0.05 % and 0.25 %.

The test speed was 20 mm/min, the clamping length 50 mm and the preload 0.1 N. The Young’s modulus was determined at a speed of 0.5 mm/min with an elongation between 0.05 % and 0.25 %. Images of cross-section of the fracture region of the tensile strength specimens were taken with a Keyence VHX-1000 microscope equipped with a VH-Z20R objective, in order to determine the fracture area for the calculation of the mechanical properties.792(cid:53)(cid:72)(cid:86)(cid:88)(cid:79)(cid:87)(cid:86)(cid:3)From the 600 g of raw material, 450 g of rounded product were obtained in the separator unit of the downer. This corresponds to a yield of 75 % in mass. The product was contaminated with large aggregates and agglomerates of particles, produced due to collisions between two or more particles during the rounding process. These agglomerates would affect negatively the packing density of powder.

These agglomerates would affect negatively the packing density of powder. For that reason, the product was sieved by using sieve-wide of 200 μm, which correspond to the x90,3 of the PSD of the feed material (see Figure 1). After sieving, a total mass of 380 g was obtained which is 63.3% of the fed mass of raw material.Figures 1a and 1b show SEM images and Figures 1c and 1d show microscope images of the raw and rounded materials, respectively. The raw material is characterized by irregular shaped particles with a broad distribution of aspect ratios, sharp edges and relative rough surfaces. In the rounded product, the roughness of the particle surface decreased considerably and the edges were rounded. A complete spheroidization was only achieved for particles smaller than 30 μm. This fact indicates that particles smaller than 30 μm melted completely and for that form completely spherical drops.

A complete spheroidization was only achieved for particles smaller than 30 μm. This fact indicates that particles smaller than 30 μm melted completely and for that form completely spherical drops. For particles sizes between 30 μm and 200μm, the process of shape transformation to spheres was incomplete, and only the region close to the surface of the particles was affected by the process. Particles larger than 200 μm were not modified after the rounding process, probably due to the very short residence time in the downer. Agglomeration as well as coalescence between two or more rounded particles can be observed. This explains the shift of the PSD of the powder towards higher values after rounding.Raw MaterialRounded Materiala)b)M E Sc)d)y p o c s o r c i mt h g i LFigure 1.

This explains the shift of the PSD of the powder towards higher values after rounding.Raw MaterialRounded Materiala)b)M E Sc)d)y p o c s o r c i mt h g i LFigure 1. Scanning electron and light microscope images of the raw and rounded powders.The volume and number-averaged particle size distributions (PSDs) of the raw and rounded powders determined by laser diffraction and light microcopy are shown in Figure 2. The volume and number-averaged PSDs of the raw material, determined by laser diffraction (Figure 1a and 1b, respectively) reveal a broad bi-modal distribution with modes at 20 μm and 100 μm respectively. The mean volume and number-averaged particle size793distribution of the raw material x50,3 and x50,0 are 120 μm and 20.9 μm, respectively. After rounding the particle size distribution was shifted towards coarse fractions due to agglomeration. The first mode remained constant, while the second mode increased to 120 μm.

After rounding the particle size distribution was shifted towards coarse fractions due to agglomeration. The first mode remained constant, while the second mode increased to 120 μm. Accordingly, the volume-averaged mean particle size increases to 132 μm while the number-averaged mean particle size remains unchanged. These facts indicate that the small particles predominantly agglomerate to the larger particles.Volume-averaged particle size distribution 0.010 1.0;1~· ,;1~· ~a)0.80.008Number-averaged particle size distribution 1.0r~v t ~b)0.070.80.06n o i t c a r f f i d r e s a L(")a0.60.40.20.0'Y/ r : Ji • f~ .. ,· Ii: \:; ·"' . J,' /J.· ~ • rJ ~/, a. "'u J/. .... .Raw Material•-· Rounded'. ..........

Volume-averaged (a and c) and number-averaged (b and d) particle size distributions of theraw and rounded powders determined by laser diffraction (a and b) and light microscopy (b and d).The number-averaged PSDs determined by light microscopy (Figure 2d) present a mode at approximately 25 μm for both, the raw and rounded powders. The number-averaged mean particle sizes of the raw and rounded powders are 30.4 μm and 34.5 μm, respectively. In the case of the volume-averaged particle size distributions (Figure 2c), no clear modes could be identified, as the image analysis resulted in several peaks. The volume averaged mean particle sizes of the raw and rounded powders are 133 μm and 143.5 μm, respectively. The results also suggest that the particle size of the raw material was slightly shifted towards higher values after rounding as consequence of agglomeration. The distributions obtained by image analysis slightly overestimate the values obtained by laser diffraction.

The distributions obtained by image analysis slightly overestimate the values obtained by laser diffraction. However, the results obtained by the two techniques are in good agreement.The number (Figure 3a) and volume-averaged (Figure 3b) particle circularity distributions of the raw and rounded powders are shown in Figure 3. The circularity distributions of the raw material reveals a broad distribution of particle shapes which agrees with the images depicted in Figure 1. The number and volume- averaged mean circularity of the raw material are 0.7 and 0.65, respectively. After rounding, the number-averaged circularity distribution was shifted considerable towards higher values while the volume-averaged was only slightly affected.

....... ---- Raw Material - -+- - Roundeda), '0.00.2...... ~J...-"'T"""--r-----t 1.0 0.40.60.8I 1.0 'T"""i"""""!!!!!!!!!!l!!!!!!!!!!!l!!!!!!!!!!!I!!!!!~-----~--, 'cv:, Cf) C 0.8 ,Q :5 ..0 ·;;:: 0.6 _. Cl)- - Raw Material - -+- - Roundedb)i5 Q) 0.4 > :.;::::; co :j E E :J 0 0.00.20.00.20.40.60.8Circularity/-Circularity /-Figure 3. Number (a) and volume-averaged (b) circularity distributions of the raw and rounded powders.The mean circularity of the particles between 0 and 60 μm increases considerably after rounding. However, for particles sizes between 60 μm and 90 μm the mean circularity remains unchanged and for particles larger than 90 μm the circularity decreases after the rounding process. The reduced change of shape was due to the fact, that large particles only melt completely at the surface, such that the shape of the particles was not highly influenced by the process.

The reduced change of shape was due to the fact, that large particles only melt completely at the surface, such that the shape of the particles was not highly influenced by the process. The formation of agglomerates during the rounding process, causing structures with low circularity (see Figure 1), contributed to a reduced rounding effect. Control of agglomeration and shape transformation require a more in-depth look into the occurring micro-processes and is the topic of separate, on- going work.0.90C cu Q)0 .85I E --"O ~0.80 Q) LO g' CJ) 0.75 .._ >, Q) > cu 0 .70 cu - ::J .._ I (.) 0 .65 Q) .._ ..0 ·-E o 0.60 ::J z;!:=.A■.A■■ Raw Material .A Rounded.A ■■ .A■.A■.A0 .55(Y'?>Cl,'oCl,r~Cl,._'J,,Cl'?>Cl9JCl' \ '],,Cl' Particle size /µm'o()\ (;)Cl\ 'o()\ (;)Cl'Figure 4.

Number-average mean circularity as a function of particle size.After rounding, the packing density of the powder remains unaltered at a value of 0.340 ±0.002 g/cm3 indicating none or only a minor effect of the rounding process on the packing density. The flowability (flow function ffc) of the raw and rounded powders are depicted in Figure 5. The flowability of the powder was improved at all the tested consolidation stresses. At low consolidation stresses (1450 Pa) the flow function increases from 1.5 to 1.8, which correspond to an increase of 21%. The improved flowability is attributed to the rounded particle edges obtained after rounding. Rounded edges would reduce the number of contact points between the particles, thus reducing the friction and resistance to flow [23,24].7951.0(.)b ..c ...... 0) C ,._ Q) ...... -o 1000 Q)Cf)">. "'O Q) C i.i= C 0 (.) C :J200040006000Major consolidation stress cr 1 / PaFigure 5.

0) C ,._ Q) ...... -o 1000 Q)Cf)">. "'O Q) C i.i= C 0 (.) C :J200040006000Major consolidation stress cr 1 / PaFigure 5. Influence of particle rounding process on flowability.Figures 6a and 6b present the images of single-layer and tensile test specimens produced from the raw and rounded polypropylene powders. Single-layers produced from the rounded powders resulted in considerably less defects and lower porosity in comparison to the single-layers from the raw materials. The rounding process also leads to test specimens with higher surface quality (lower surface roughness) based on qualitative observation.a)b)Figure 6. Influence of rounding process on a) defects of single-layer specimens and b) surface quality oftensile specimens.The results of the tensile strength test on the produced samples are depicted in Figure 7. Both, the Young´s modulus and the stress at break increased pronouncedly after rounding.

10 ~ co C ..Cl 8 ....... 2 .Q ....... co co Cl) O> Cl) C Q) .._ 0 ....... (f) w642000PP rawPP roundedFigure 7. Tensile properties of SLS produced test specimens.ConclusionsIn this work, we presented an experimental study on the rounding of commercial polypropylene powders using a modified downer reactor. The process result in a slight shift of the particle size distribution towards higher particle sizes as consequence of agglomeration. The rounding process mainly affected the surface of the particles rather than the shape: Produced particles show a shape like the feed, however, with rounded edges. The modification of the surface of the particles contributed to an improvement of the flowability of the powder.

fect of Particle size on SLS and post-processing of Alumina with Polymer bindersKamatchi Subramanian, l.W. Barlow and H.L. MarcusCenter for Materials Science and Engineering The University of Texas at Austin.AbstractAlumina powders with mean particle size of 15 J.lm, 2 J.lm and 0.5 J.lm were processed by SLS using polymer binders. The 2 J.lffi and 0.5 J.lm powders were given a thermal agglomeration treatment before SLS. The green shapes after SLS were infiltrated with a colloid of aluminum oxide. After infiltration the samples were given suitable heat-treatment to remove the polymer binder by thermal decomposition followed by a sintering treatment at 16OOC. Green densities were in the range of 45% of theoretical density for the agglomerates of 2 Jlm and 0.5 J.lffi powders while it was about 36% of theoretical density for the 15 J.lffi powders.

Green densities were in the range of 45% of theoretical density for the agglomerates of 2 Jlm and 0.5 J.lffi powders while it was about 36% of theoretical density for the 15 J.lffi powders. Sintered densities were about 55% of theoretical density for the samples from agglomerates of 2 J.lffi and 0.5 Jlm powders while it was 42% of theoretical density for the samples from 15 J.lm powders. The strength of samples were measured in the green state and after sintering to determine the effect of particle size.IntroductionSelective Laser Sintering (SLS) of ceramics using polymer binders has been investigated at the University of Texas [1-5]. Single phase alumina has been obtained by processing the ceramic powder as follows [6]: i) SLS with polymer binder, ii) infiltration of the green shape with alumina colloid, iii) removal of the binder by thermal decomposition and iv) sintering at high temperatures of the ceramic samples after binder removal.

In this paper we will examine the effects of particle size on the SLS and post-processirtg of alumina using polymer binders.Experimental workAlumina powders of 15 Jlm mean particle size were spray-coated with a copolymer containing 80 mole% methyl methacrylate and 20 mole% butyl methacrylate. The 2 Jlm alumina powders could not be processed by spray-drying with polymer binders. They were given a thermal agglomeration treatment at 14000c for 3 hrs [5]. After the agglomeration treatment the cake that formed was broken down and agglomerates less than 125 Jlm were mixed with the polymer powder (20 vol% polymer) obtained by spray-drying of polymer emulsion. For the 0.5 Jlm alumina powder the thermal agglomeration treatment was conducted at 13000C because the cakes obtained at 14000C could not be broken down easily due to the accelerated sintering of the small particle size.

These agglomerates were also mixed with the polymer powder (20 vol% polymer) obtained by spray-drying.The spray-coated powder and the powder mixtures were processed by SLS and bend strength specimens (3 I xl l xO.25") were made. The specimens in the green state (as SLS) were tested for density and green strength. Some of the specimens were infiltrated with alumina colloid346by immersing the samples in the colloid. After inftltration with colloid the samples were dried at 120°C and then given a debinding treatment [6]. This was followed by a sintering treatment at 16000 C for 6 hrs.Results and DiscussionIn the case of green samples the particle size affects the green strength depending upon the surface area to be bonded by the polymer.

A comparison of the strengths of green samples from agglomerates of 0.5 Jlm alumina mixed with 20 vol% copolymer powder, agglomerates of 2 Jlm alumina mixed with 20 vol% copolymer powder and 15 Jlm alumina coated with 25 vol% copolymer is shown in Figure 1.C':& ~ ~ ..= ...... .. eJ:) = 1.52.0~ ...... c:.f} "0 1.0 = =~0.500.5 J.UIl2J.UIl A- - - - A 0I6 - A AA00• - AA00AA0.0 01234561Energy density, cals/cmzFigure 1. Effect of particle size on green strength of samplesFor the 0.5 Jlm alumina the agglomeration into spherical agglomerates at the lower temperature probably resulted in weakened agglomerates and therefore lower bend strength. The 15 Jlm alumina has a higher volume fraction of polymer. However it will have more area to be wet by the polymer than the 100 Jlm agglomerates made from the 2 Jlm alumina. Hence the strength obtained is lower even with higher volume fraction of the polymer.

Hence the strength obtained is lower even with higher volume fraction of the polymer. The decrease in bend strengths at higher energy densities is associated with polymer degradation and is reported elsewhere [6,7]. The densities of the samples are shown in Figure 2. Fully dense alumina is 3.97 glcm3.From these it can be seen that the density for the samples from agglomerates of 2 Jlm powder marginally increases with energy density, and the density of samples from this mixture is slightly higher than that obtained from agglomerates of 0.5 Jlm. The difference in green strengths cannot be related to the difference in density between the two materials. The difference in strength347is probably due to the difference in surface area to be wet by the polymer. The density of the 15 Jlm powder samples is lower than that of the other two because that powder does not pack as well as the other two and the polYmer content is higher.e~<:J ........

This is due to the highersintering kinetics of the smaller particles at the sintering temperature. The independence withenergy density is due to the high temperature sintering controlling the strength instead of thepolymer distribution.Regarding the strength of samples after sintering, Figure 4, a couple of general observations can be made. The samples from agglomerates of 0.5 Jlm powder are stronger thansamples from other two because of the higher sintering kinetics for the smaller particles. Thestrengths all decrease with increasing incident energy used during laser processing of the green samples. With increasing incident energy density more of the polymer is decomposed. The rate ofheating in SLS processing is of the order of 6000K/10ms [8]. This high heating rate will cause sudden expansion of gases formed by decomposing the polymer, in the powder bed leaving behind larger defects.

Effect of particle size on strength of sintered specimens.349Figure 5. Cross-section through sintered sample from 2 Jim alumina [6].Figure 6. Cross·-section through sintered sample from 0.5 Jim alumina350the strength is inversely proportional to defect size, the strength decreases with increasing energy density. Due to the difficulty of evaluating fracture surfaces of porous materials, this hypothesis could not be confinned.Microstructure of sintered sPeCirnensIt may be seen that there are lacy regions corresponding to alumina from sintered colloid. In the case of the 2 J,lm samples the agglomerates remain spherical even after the 1600 OC heat treatment. However in the 0.5 J,lm sample the agglomerates no longer appear spherical because of extensive solid state sintering and wide neck formation during sintering.

However in the 0.5 J,lm sample the agglomerates no longer appear spherical because of extensive solid state sintering and wide neck formation during sintering. This contributes to the higher strength observed.The section through sintered samples are shown in Figures 5 and 6.SummaryThe effect of particle size on SLS and post-processing of alumina has been examined using three particle sizes of alumina 15 J,lm, 2 J,lm and 0.5 J,lm. The finer powders needed to be agglomerated before processing by SLS. After SLS processing the samples undergo infiltration, debinding and sintering at 1600 OC. The sintered densities obtained were about 55% of theoretical density for the 0.5 J,lm and 2 fJm powders while it was about 40% for the 15 J,lm powders. Sintered samples from the agglomerates of 0.5 J,lm were stronger than the other two because of the higher sintering kinetics of the finer particles.

Sintered samples from the agglomerates of 0.5 J,lm were stronger than the other two because of the higher sintering kinetics of the finer particles. The sintered samples exhibited effects due to the decomposition of the polymer during the SLS process.AcknowledgmentSupport for the research ~eported here was received from ONR Grant # NOoo14-92-J-1514.References1. Vail, N.K., and Barlow, J.W., "Microencapsulation of Finely divided Ceramic powders", Proceedings of the Solid Freeform Fabrication symposium, August 6-8, 1990, editors: J.J. Beaman,H.L. Marcus, lW. Barlow and D.L. Bourell, The University of Texas at Austin, p.82. Vail, N.K., and Barlow, J.W., "Effect of Polymer coatings as intermediate binders on Sintering of Ceramic parts", Proceedings of the Solid Freeform Fabrication symposium, August 12-14, 1991, editors: H.L. Marcus, J.J. Beaman, J.W. Barlow, D.L. Bourell and R.H. Crawford, The University of Texas at Austin, p.1953.

Marcus, J.J. Beaman, J.W. Barlow, D.L. Bourell and R.H. Crawford, The University of Texas at Austin, p.1953. Vail, N.K., Barlow, J.W., Beaman, J.1., Marcus, H.L., and Bourell, D.L., "Development of a Poly(Methyl Methacrylate-co-n-Butyl Methacrylate) Copolymer Binder system," Journal of Applied Polymer Science, 1994, 52, p. 7893514. Subramanian, P.K., Zong G., Vail N.K., Barlow, I.W., and Marcus, H.L., "Selective Laser Sintering of Alumina", Proceedings of the Solid Freeform Fabrication symposium, August 9 editors: H.L. Marcus, 1.1. Beaman, J.W. Barlow, D.L. Bourell and R.H. 11, 1993, Crawford, The University of Texas at Austin, p. 3505. Subramanian, P.K., Vail, N.K., Barlow, I.W., and Marcus, H.L., "Anisotropy in Properties of Alumina processed by SLS", Proceedings of the Solid Freeform Fabrication symposium, August 8-10, 1994, editors: H.L. Marcus, 1.J. Beaman, J.W. Barlow, D.L. Bourell and R.H. Crawford, The University of Texas at Austin, p. 3306.

Research at the IRCCYN laboratory (France) focus on the impact of cladding or welding head programming on dimensional properties and microstructure of the manufactured parts. In this context, path strategies are studied and different strategies are proposed and compared. This paper presents results for welding path strategies. The integration of deposition head on a 6 axis robot system was used to improve slicing strategies and parts.I.IntroductionOne of the most important topics of research leads by the MO2P Team (Modelisation and Optimization of Process of Production, IRCCYN France) on hybrid modular tools is to choose between different processes the best one to manufacture modules [1-2]. A way to have a large panel of processes capable is to qualify emergent processes. Additive processes are now a great tool to manufacture directly metallic parts with high technicality.

A way to have a large panel of processes capable is to qualify emergent processes. Additive processes are now a great tool to manufacture directly metallic parts with high technicality. Technologies like Laser Engineering Net Shaping (LENS), Direct Metal Deposition (DMD), Electron Beam Melting (EBM) or Gas Metal Arc Welding (GMAW) are able to built fully dense parts with great mechanical skills. Additive processes are currently under improvement. T.Wohlers showed that trends in process improvement focus upon material response to additive fabrication, powder-flow studies, tolerance, and machine automation through feedback control. Other process improvements include the design and manufacture of deployable structure, residual stress analysis, part quality and process optimization [3]. Rapid Manufacturing by Gas Metal Arc Welding uses metal fused by a welding torch to manufacture fully dense parts in layers.

Rapid Manufacturing by Gas Metal Arc Welding uses metal fused by a welding torch to manufacture fully dense parts in layers. These parts have excellent mechanical properties and their structure was fully dense with minimal inclusions [4]. GMAW can be performed using readily available equipment which has been investigated and is well known. A review of previous works on rapid manufacturing by robotic GMAW showed the main disadvantages are high temperatures of substrates which create warpage and residual stresses, dimensional accuracy and surface quality, weld defects and unfilled voids, difficulty to control the process and geometric instability [5]. Research on this domain shows that precision and thermal conditions are deteriorate when the number of layers grows up. For P.M.Dickens, “inaccuracies in the welding and robot parameters can cause cumulative errors, resulting in the torch being too close or too far away from the surface” [6].

For P.M.Dickens, “inaccuracies in the welding and robot parameters can cause cumulative errors, resulting in the torch being too close or too far away from the surface” [6]. I.Kmecko found that geometrical deterioration is due to small defects in the previous layer that become more and more amplified in the subsequent layers [7]. One way to solve geometrical and thermal instability could be to combine GMAW and CNC machining [8-10]. For Z.Jandric, the principal disadvantage of this method is its slow speed. For him, if the surface of each deposited layer would be smooth enough, layers could be deposited without introducing the milling operation [11].Another way to overcome geometrical and thermal problems is to study the effect of weld settings [4], or the effect of weld path strategy [5] on the object geometry and mechanical skills. The aim of this paper is to focus on choice of path strategies for GMAW to optimize flatness of parts build by this way.II.

The aim of this paper is to focus on choice of path strategies for GMAW to optimize flatness of parts build by this way.II. Manufacturing methodology Starting with CAD model, the first step of the methodology used in this study is to choose the best process by a geometrical analysis. Then, if an additive process is selected (Fig.1), the second step consist in determination of the substrate form. It could be a base plate or an existing part. The use of a mechanically welded or sheet incremental formed preform [14-15] as a substrate is a possible way to build part without machining process. Moreover, these substrate forms will be obtained without tooling. The substrate form induces the choice of a slicing strategy.

Moreover, these substrate forms will be obtained without tooling. The substrate form induces the choice of a slicing strategy. The next step is to find a path strategy and a parameter set that could manufacture part’s geometry.Fig.1: Methodology applied to additive processesThe main objective of this methodology is to find the best way to obtain a fully dense part with good dimensions and a great accuracy. Applying to metal deposition of a rectangular shape part, Gas Metal Arc Welding process has been chosen. A base plate is the best substrate geometry in this case. (Fig.2).Fig.2: A test part composed by 4 layers with rectangularConcerning the slicing strategy, the question is to know the best way to obtain a good final geometry.

(Fig.2).Fig.2: A test part composed by 4 layers with rectangularConcerning the slicing strategy, the question is to know the best way to obtain a good final geometry. Two strategies will be compared: 3 axis strategy, which consist in stacking planar layer as copy of the final surface and irregular slicing strategy which consist in deposit layers with different forms or thickness.III.Experimental approachTo have a geometrically stable process, it is important to find a strategy that allows us to obtain a surface flat enough to be a good base for the next layer or a strategy that can cancel the amplification of the deterioration of surfaces. A review of previous works leads in welding techniques to manufacture wide parts shown that a technique called layer filling constrained strategies, like “double spiral overlay welding” [12] allow to limit the effect of thermal stresses. These techniques consist in deposit welds in layers with two sets of weld settings.

These techniques consist in deposit welds in layers with two sets of weld settings. The first set of welds, called ridge welds [5], is used to deposit welds that will constrain the second set. Then, the second one, called through welds [5], is use to fill the space left between ridge welds already deposit. Ridge welds constraint through weld’s melt pool (Fig.3).Fig.3: Concept of layer filling constrained strategyThen, it is possible to let cooling times between ridge and through weld deposition. It is a good technique for the process stability, because the ridge welds are all deposited with the same constraint, and the through welds too.

It is a good technique for the process stability, because the ridge welds are all deposited with the same constraint, and the through welds too. Coupled with a contour strategy (Fig.4), layer filling constrained technique is geometrically stable and had a good thermal stability due to the most cooling stops [5].: Through weld : Ridge weldFig.4: Layer filling constrained strategy coupled with Contour strategyM.Siminski had shown that “self-constrained” path strategies were fairly flat overall except that they had periodic wave-like shape coinciding with the ridge and through welds. It could be seen that was caused by the through welds being a little too high for the ridge welds, due to imperfectly matched weld settings.”[5].

It could be seen that was caused by the through welds being a little too high for the ridge welds, due to imperfectly matched weld settings.”[5]. The first trial made with a GMAW robot confirm this observation (Fig 5).Fig.5: Cartography of wave like effect on first trialIt is important to control this “wave-like shape” by testing different parameters sets to obtain a great level for flatness of the final surface manufactured. Another way to obtain flat surface has been tested.

Another way to obtain flat surface has been tested. It consists in alternating or not through and ridge position between each layer (Fig.6).Layer 2Layer 3Layer 4: Through weld : Ridge weldFig.6: Ridge and through welds inversionIV.Experimental setup1) DescriptionFirstly, a series of solid object weld trials was carried out with a robotic GMAW in order to find more influent parameters of a self constrained contour strategy on:(cid:129) Flatness (cid:129) Parallelism (cid:129) HeightThen, in a second time, the choice was made to focus on optimization of this flatness. The test part selected to test the flatness of layers built by robotic GMAW consisted in a cube composed by 4 layers with rectangular shape. All the experiments was made with a constant parameters set for ridge welds in order to work only on through welds parameters (Tab.1). Due to experiment constraint, like control between two layers deposition, cooling time was fixed.

Due to experiment constraint, like control between two layers deposition, cooling time was fixed. The wire diameter was fixed to 1.2mm.Ridge weldsVoltage (V) 22Wire feed rate (m/min) 5.3Travel Speed (cm/s) 0.4Tab.1a: Ridge welds parametersThrough welds space between ridge welds (mm) Voltage (V) Travel Speed (cm/s)10 to 1524 to 260.2 to 0.6Tab.1b: Through welds parameters setsAs the effect of the space between ridge welds on flatness had been tested, length and width accuracy of the process in this paper was not studied. The dimensions of parts manufactured varied with settings. Parallelism and height were controlled only to validate the result of trials.2) Welding cell and equipment usedThe welding robot used for this experiment was a KUKA KRC1 Robot coupled with a Fronius Trans Synergic 4000 welding power supply (Fig.7a).

Controls of flatness, parallelism and height were all been made with an ATOS sensor (Fig.7b).Fig.7a: KUKA KRC1 RobotFig.7b: ATOS STD sensorThis flexible optical measuring machine is based on the principle of triangulation. Projected fringe patterns are observed with two cameras. 3D coordinates for each camera pixel are calculated with high precision, a polygon mesh of the object’s surface is generated. With this sensor, an accurate 3D mapping of manufactured surface had been made. Then, the soft create least squares plane and compute distances between points and this plane to obtain flatness defect [13].V. 1) Surface topography of the first layerResultsTwenty three parts had been manufactured with different parameters set to identify some topography types. A choice has been made between those trials to find the most significant parts manufactured that could be the base of the second step of the experiment.

A choice has been made between those trials to find the most significant parts manufactured that could be the base of the second step of the experiment. Parts chosen must had repeatable results, a regular shape and no welding defect.The results of surface analysis of the first layer showed two types of surface topography (Fig.8):(cid:129) Flat surfaces (cid:129)“Wave like shape” surfacesFig.8a: Flat surface topography (first layer)Fig.8b: “Wave like” surface topography (first layer)Measurement showed a flatness defect value of approximately 1.4mm for flat surfaces and 2.4mm for “wave like shape” samples. But, with the observation of the surface quality, the meltpool size of parts with flat layers grow during weld deposition.

But, with the observation of the surface quality, the meltpool size of parts with flat layers grow during weld deposition. For this reason, a collapse of this type of surface topology parts is predictable when the number of layers will grow up.2) Surface topography of the fourth layerThe next step of the experiment was to test stability of flatness in the subsequent layers. So, the comparison of the use or not of the“Ridge and through welds inversion” strategy on the two surfaces topology was tested. Cartographies of surfaces obtained highlight the growth of the flatness defect for “flat surface topology” parts. The collapse of parts is trivial and appears first with the inversion of weld parameters. This new strategy was not adapted to flat layers. The fabrication was stopped after the second layer for this strategy because sample’s dimensions were no more respected. The other strategy was conducted to the end.

The fabrication was stopped after the second layer for this strategy because sample’s dimensions were no more respected. The other strategy was conducted to the end. (Fig 9)Fig.9: Collapse of “flat” surface topography (fourth layer)The collapses on “flat” surfaces were due to the weld pool size at the end of the program. The weld pool of the “wave like shape” samples had a constant size. It explains the better stability of those parts. For “wave like shape” surface topology parts, results show that flatness defect grows less with the new strategy than with the repetition of same layers.

For “wave like shape” surface topology parts, results show that flatness defect grows less with the new strategy than with the repetition of same layers. (Fig.10)Fig.10a: Flatness defect without “Ridge and through welds inversion” strategy (fourth layer)Fig.10b: Flatness defect with “Ridge and through welds inversion” strategy (fourth layer)Flatness defect (in mm) Layer n° Flat surface without “Ridge and through welds inversion” strategy (Strat.1) Flat surface with “Ridge and through welds inversion” strategy (Strat.2) “Wave like shape” surface without “Ridge and through welds inversion” strategy (Strat.3) “Wave like shape” surface with “Ridge and through welds inversion” strategy (Strat.4)Flatness defect6554,4m m n43,5Strat.1it c e f e D32,4 2,4Strat.3Strat.421,410012345LayersFig.11: Flatness defectOn each case, higher points are located at corners of each weld.

Deeper points of these two cartography are located at the “start and stop” point of each weld. Some method has been developed in literature to solve this problem [16]. It will not be treated here. Using the “Ridge and through welds inversion” strategy, Pyramid effect (central area is higher than contour area) can be identified (Fig.10b). It was due to the lack of space let by the first weld to fill the central zone. This could be solved by modified program of through welds. Periphery area had a better flatness than the periphery area of samples which not used this strategy. Height of welds was more regular too. In fact, on parts build without this strategy, there is an offset of welds position due to amplification of defects which caused height variation of welds (fig.12).Fig.12: Offset of weldsThe “wave like shape” effect is more important, the distance between wave through and wave crest is higher.

Flatness measurements show that defects grow with the number of layers, but grow less with “ridge and through welds inversion” strategy than without. In this case, the use of this strategy is efficient.1 1.4 1.4 2.4 2.44 5 - 4.4 4VI.ConclusionThe “Ridge and through welds inversion” strategy has been applied with success. It reduced the growth of the flatness defect (Fig 11). The interest of this result is that there is no need to start with a perfectly flat surface to obtain a perfectly flat surface at the end of an additive manufacturing process. It highlights the significance of the geometrical study of parts to choose the best slice strategy and path strategy for the manufacturing method developed in this paper. Then, starting with the irregular slicing strategy, it will be interesting to develop a “top to bottom” slicing strategy instead of a classical “upward” strategy (from substrate to skin).

Then, starting with the irregular slicing strategy, it will be interesting to develop a “top to bottom” slicing strategy instead of a classical “upward” strategy (from substrate to skin). The surface cartography is a great tool to obtain flatness information and it permits to know the exact height of points of the part. Then, it is possible to guaranty tool/parts distance in the next layer program. The next step of this study is to optimize parameters set to reduce (or cancel) the defect value and defect growth. The study had been made on simple parts with well known available equipment.

The next step of this study is to optimize parameters set to reduce (or cancel) the defect value and defect growth. The study had been made on simple parts with well known available equipment. It will be interesting to develop irregular slicing strategy with another process, like Laser Engineering Net Shaping (LENS), on more complex geometries, with other substrate forms, as described in manufacturing methodology.REFERENCES [1] Mognol, P., Rivette, M., Jegou, L., Lesprier, T., 2007, A first approach to choose between HSM, EDM and DMLS processes in Hybrid Rapid Tooling, in Rapid Prototyping Journal, Vol 13 pp 7-16. [2] Rivette, M., Hascoët, J.Y., Mognol, P., 2007, A based graph methodology for hybrid rapid design, in Journal of Engineering Manufacture, Vol 221, N°4, pp 685-697. [3] Wholers, T., 2007, Rapid prototyping, tooling and manufacturing state of the industry, Wohlers report.

fect of Polymer Coatings as Intermediate Binders on Sintering of Ceramic PartsN. K. Vail and J. W. Barlow Department of Chemical Engineering The University of Texas at AustinAbstractIt has been proposed that a thin polymer layer applied to the surfaces of finely divided ceramic powders would serve as a suitable intermediate binder for Selective Layer Sintering of ceramic pans. In this nd particle size study, the effects of completeness and fraction of coa ridized glass are distribution on sintering rates and strengths of coated examined. The effect of the coating as a binder during Selective Layer Sintering as well as the binder's burnout capability during post processing steps are also investigated. (Key Words: Microencapsulation, Glass, Sintering, Polymer, Spray Drying)IntroductionSelective Layer Sintering (SIS) is an innovative process for producing three dimensional objects directly from a CAD database.

SIS has found use in the rapid prototyping of commercial parts using polymer and wax materials. However, parts made of these materials have limited uses, such as models and patterns for investment casting. Methods to extend the capabilities of SIS to the manufacture of ceramic and metal parts is the focus of several research projects.In a previous study (Vail and Barlow, 1990), spray drying was shown to be an effective and efficient method to encapsulate alumina with a polymer binder. In this ongoing study, spray dryin is used to encapsulate spheridized glass particles. Green parts were made from coa glass materials using SIS. These green parts were post-processed in a furnace to observe the effects of binder burnout and to observe dimensional and density changes in the sintered part.Materials and MethodsSamples of UCAR 430 Acrylic Polymer Latex were obtained from Union Carbide Corporation.

UCAR 430 is a styrene-methylmet .tylacrylate-methacrylic acid copolymer with a glass transition of approximately Spheridized soda-lime glass was 'dized soda-lime glass, obtained in two obtained from Potters Industries, Inc. The s different particle size distributions, are des' 5000 and A-3000 glass. These two size distributions have mean particle sizes of 5.0±2.5IJ.m and 24.0±15IJ.m, respectively, as measured by a Coulter Multisizer.Glass powder was coated with polymer by frrst mixing the glass powder with the polymer latex to produce a slurry. This slurry was then spray dried using an Anhydro Laboratory Spray Drier #1. This drier has a cylindrical drying chamber with conical base. The drier is equipped with a Ihp air blower which provides positive pressure drying air to the chamber. Drying air can be heated from 125°C to 300°C via a 9kW electrical heater.

The drier is equipped with a Ihp air blower which provides positive pressure drying air to the chamber. Drying air can be heated from 125°C to 300°C via a 9kW electrical heater. Feed to the drier can be atomized either by centrifugal or by pressurized nozzle means.195Centrifugal atomization was used for the purposes of this paper. The centrifugal atomizer contains a four veined atomizing wheel which is driven by a motor to speeds of 50000 revolutions per minute. The atomizer wheel speed was verified by an optical tachometer. Dried product powder is separated from the drYing air by a cyclone.Characterization of Spray DrierSeveral experiments were conducted to determine the effects of inlet drYing temperature, solids content through the drier, and polymer content of the product on the product particle size. Ten samples each were run at two inlet drying temperatures, 1300C and 160°C. At each temperature, samples contained 25% and 50% total solids by weight.

Ten samples each were run at two inlet drying temperatures, 1300C and 160°C. At each temperature, samples contained 25% and 50% total solids by weight. Each total solids concentration was further divided into 1%, 2.5%, 5%, 10%, and 15% polymer weight of total solids for a total of twenty samples. Each sample contained a fixed amount of A-5000 glass spheres. Other than water being added to the mixture of materials to attain desired concentrations the materials were used as supplied. Typical sample compositions are listed in Table 1. Spray drier operating conditions for selected materials are shown in Table 2.Glass, (g)250Table 1. Spray drier sample compositions Polymer Content, (wt. % total solidS) 1.0 2.5 5.0 10.0 15.0 1.0 2.5 5.0 10.0 15.0Total Solids, (wt %)UCAR 430 Emulsion, (~) 5.7 14.6 29.2 61.7 100.3 5.7 14.6 29.2 61.7 100.325.050.0Water, (~)754.4 761.4 I 773.4 799.4 826.2 249.3 248.6 247.1 243.8 238.0.IITable 2.

Spra' drier operatin J conditions s~~:-nn=======Ma===te==n=='a==1=====P}~).In1e~o~mp.,Outlet Temp., (90)UCAR430 Emulsion Geon 151 Emulsion GlasslUCAR 430 Mixture (25%/50% Solids) GlasslUCAR 430 Mixture (25%/50% Solids)9012080/90100/1103750037500375003750013016013016060636065196Particle Size and Polymer Coating AnalysisParticle size distributions of glass materials and of all spray dried materials weredetermined by Coulter Multisizer measurements. Polymer coatings were observed by scanning electron microscopy and polymer content of the spray dried powders were determined by Thermal Gravimetric Analysis (TGA).Selective Layer SinteringThree samples of A-5000 glass were processed via spray drying to produce coatedparticles with polymer contents of 5%, 10%, and 15% by weight.

Pure DCAR 430 polymer was spray dried for the purposes of mixing with the same spherical glass to produce samples with comparable polymer compositions to the coated samples. The prepared samples were processed by SIS using a CO2 laser operating at three different scan speeds. All part processing was conducted at ambient bed temperatures and in an open atmosphere. Twelve parts, one inch square by one-quarter inch thick, were made from each coated glass sample. SIS processing conditions are presented in Table 3.Table 3. Selective Layer Sintering operating conditionsPower, (W)25Layer Thi kn cess, (It) 0.003Scan Speed, (ips)32.0, 40.0, 64.0Roller SpeedsLinear (ips) , 4.0Rotational, (rpm) 82.0Post-Processing of Selective Layer Sintered Parts.To examine effects of post-processing on SLS parts, the samples previously prepared were heat treated using a high temperature box furnace.

Samples were removed from the SLS machine, brushed to remove loose powder, weighed and measured, and packed in fine alumina in a porcelain crucible. Samples were placed in the center of the furnace and brought to temperature at a heating rate of lOOC/min. As temperature, the polymer was burned off. Furnace the sample came to sinte yaK-type thermocouple placed directly above the sample temperature was monitore crucible. The thermocouple was connected to a digital thermometer. All samples were processed at 637°C.Results and DiscussionScanning electron micrographs of Potters A-5000 soda-lime spherical glass and a es 1 and 2, erical and of particles to betypical spray dried coated sample of the same glass are shown in rial to be near respectively. The pure glass micrograph shows th the spray almost uniform size. The coated glass micrograph r of polymer. agglomerates of the spherical glass coated with at' sample is presented in Figure 3.

Spray dried particle size results199Figure 4. Particle. size is influenced both by the·.•amount.of solids in the•. spray drier feed and by the amount. of polymer.inthe product. Temperature. dicinot .appear to playa role in product particleisize. As expected, particle size increases/with<increasingpo er content in. the product and,Results of particle .size.measurements arediscl.lssedbyMasters (1972) who statedthat there are two reasons why larger particle sizes result at l~·trger solids loading. First, lower moisture content of tbe atomized droplets facilitate higher drying ratesresl.lltinginhollow particles and overall higher bulk densities. This .effect .• is most prevalent•• infeedscontaining dissolved .solids. Second, the initial of moisture evaporation from atomized droplet size is reducedin:r>roportion to the sizes,.•.• feedswithbigh solid the •• droplet. Therefore, assuming.

samples with polymer ""''-IJLJu• ...,...... IJJ........'......lLJlIJJ mixture of glass and polymer is1I..4"1crl11"'.::ll."..........r\1I"Q>1I"sufficient strength the single. sample of coated and mixed part made from the from polymer coatedIn fact, the green part made of mixed glass and pol~er is too delicate to handle. glass. It is for these reasons that only one sample of glass and polymer mixture was processed. In addition to producing poor quality parts with low green strength the mixed sample could only be processed at the fastest laser scan speed since significant part curling occurred at the slower scan speeds..Figures 6 and 7 show sintering results of green parts fired for times of Y2, 1, 2, 4, and 8 hours. Green parts made from 15% wt. polymer coated glass SIS processed at three laser scan speeds were sintered. Figure 6 shows the final relative densities of these parts to be independent, within experimental measurements, of the laser scan speed.

Figure 6 shows the final relative densities of these parts to be independent, within experimental measurements, of the laser scan speed. Figure 7 shows the axial shrinkages of parts at all sintering times to be very consistent in all directions. The z-direction does show a slightly higher shrinkage at all sintering times, but it is not drastically· different. This indicates the porosity in the z-direction to be equal to the porosity in the xy plane.Figure 8 shows densities of green parts made from the three different coated glass materials. To determine the significance of these green part densities, the density of the bulk coated glass powder was measured. The powder density was measured first by allowing a fixed amount of powder to settle freely in a graduated cylinder. The cylinder was then tapped vigorously to settle the powder and the density again determined. Each measurement was repeated four times.

The cylinder was then tapped vigorously to settle the powder and the density again determined. Each measurement was repeated four times. It is interesting to note that green densities lie between the densities observed for untapped powders and tapped powders. In fact, densities of all green parts produced from the three coated samples varied only 1.04% from the mean relative density of 36.0%.E r - 64.0 ips32.0 ips46Sintering Time. hi'10Figure 6. Densities of glass parts at different sintering times.201§15 r··········,··;t...;...;.;;;.1··················!···············..~.....t' "tS....P-<i 10 I-··········..·..·..f ..··;················..··+-·..·········;..;.51-···..············....+···;·······..·······..·+2 . Sinterlng Time, hr....Figure 7. Axial shrinkages of sintered glass parts.": ' ;' , ' : ;,"500'I t··········..·..·..··...l··..· ·..··,...... -u- - ntapped : I I ~::Part ...........[;1 +....... . 1................

Axial shrinkages of sintered glass parts.": ' ;' , ' : ;,"500'I t··········..·..·..··...l··..· ·..··,...... -u- - ntapped : I I ~::Part ...........[;1 +....... . 1................ ~ PowderBed ! ...- i....I0i.......... it135i ; 'o i i i i i Ie i ' " i I.'..Figure 8. DensitiesofpOlymerOO3tedglass powders.To/furtherinvestisate. theseresults,twoisupplementary sets of experilnents were In the firstisetofexp~ri)nentspowderbeddensities ofpureA,-SOOOglass and conductecj. the polymer coated A-SOOO glass were measured by determining the depth of a fixed number202of powder layers deposited by the SLS machine and then weighing the powder. Powder leveling conditions for these experiments were the same as those listed previously in Table 3. A second set of experiments was conducted to observe the effect of adding a larger size spherical glass to a sample of 15% wt. polymer coated glass on powder bed densities.

A second set of experiments was conducted to observe the effect of adding a larger size spherical glass to a sample of 15% wt. polymer coated glass on powder bed densities. Two samples were prepared by adding A-3000 glass to the coated. glass to reduce the polymer content of the entire sample to 10% and 5% wt.Figure 8 also shows the results of bed density measurements of the coated glass powders. The powder bed densities are only slightly higher than the untapped powder density and are considerably lower the tapped powder densities. Part densities lie slightly above the bed densities and green part density increases are related to the amount of polymer present. Further, powder bed density is also related to the amount of polymer present. Indeed, the presence of polymer causes a drastic change in bed density when the data are compared to the observed relative bed density of about 60% for pure A-5000 glass.Fipre 9.

Indeed, the presence of polymer causes a drastic change in bed density when the data are compared to the observed relative bed density of about 60% for pure A-5000 glass.Fipre 9. Densities of coated glass spheres mixed with large size glass spheres.Figure 9 shows the results of mixing 15% wt. polymer coated A-5000 glass with two different proportions of larger sized A-3000 glass. The bimodal powder densities are significantly higher than the densities of plain coated glass powders. Reed (1988) states these results are as expected since the smaller coated glass particles fill the interstices of the large glass compact. Why the green densities are lower than the bed densities and, in the case of 33% coated glass, lower even than the untapped density is unclear. Still, it is interesting to note that green parts from these mixtures showed sufficient green strength to permit handling.Finally, post processing was performed on these green parts.

Still, it is interesting to note that green parts from these mixtures showed sufficient green strength to permit handling.Finally, post processing was performed on these green parts. Two parts from each of the two bimodal samples were fired at 637°C for two hours. These results are compared203to results of similar post processed 15% wt. coated glass green parts in Figure 10. For the same conditions the bimodal parts show smaller density changes and therefore smaller axial shrinkages.42.51-+ )", .\.." + ~ --R--f~ 40.01-···············..·+·······················+·l\l.··+;-'1'i ~37.51-··················..,,+.··············;·····""::;"Y"'~ ,~;l4 . Weig1)t.,percentPolymer COntent8,12,1620Fipl'elCl •FireddeJl$i~esof CQate<.lglass green. parts and coated glass mixed with larger glass green parts,ReferencesVail, N.K..

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceTHE EFFECT OF POLYMER MELT RHEOLOGY ON PREDICTED DIE SWELL AND FIBER ORIENTATION IN FUSED FILAMENT FABRICATION NOZZLE FLOWZ. Wang and D.E. SmithDepartment of Mechanical Engineering, Baylor University, Waco, TX 76798AbstractShort carbon fibers suspended in the polymer feedstock enhances the mechanical performance of products produced with Fused Filament Fabrication (FFF). As the melted filament is extruded and deposited on a moving platform, the velocity gradients within the polymer melt flow orientate the fibers, and the final orientation has a direct effect on the mechanical properties of printed bead. This paper numerically simulates an FFF nozzle flow, including the extrudate material beyond the nozzle exit.

This paper numerically simulates an FFF nozzle flow, including the extrudate material beyond the nozzle exit. Finite element simulations of the extrusion process are performed with Generalized Newtonian Fluid (GNF) models and a viscoelastic rheology model, included in ANSYS Polyflow, to evaluate the polymer melt velocity field and predict die swell. Fiber orientation tensors are computed along streamlines using the Fast Exact Closure and Folgar-Tucker isotropic rotary diffusion. The predictions indicate that shear thinning behavior reduces the die swell but viscoelastic rheology significantly intensifies the extrudate swell.

The predictions indicate that shear thinning behavior reduces the die swell but viscoelastic rheology significantly intensifies the extrudate swell. Orientation tensor values calculated from the flow results of the viscoelastic model yields lower principal alignment in printed beads than those computed with GNF models.IntroductionFused Filament Fabrication (FFF), which is also known as Fused Deposition Modelling (FDM), is a widely used Additive Manufacturing (AM) technique due to its rapid manufacturing speed and low-cost. The FFF process is an extrusion-based process where polymer feedstock materials are melted and deposited on a heated platform layer-by-layer to form three-dimensional (3D) objects [1]. Recently, Oak Ridge National Laboratory (ORNL) and their associated collaborators have extended the FFF process from small scale rapid prototyping to large scale manufacturing for parts and tooling [2].

To achieve relatively high dimensional accuracy and superior mechanical performance, carbon fiber filled polymers are employed the large scale FFF system. Studies have shown that adding short carbon fibers into the neat polymer yields a composite with improved material performance [3].The prediction of fiber alignment within the printed bead is pivotal in the FFF extrusion process since fiber orientation greatly influences the material properties of solidified parts [3]. Evans, et al. [4] and Libscomb, et al. [5] showed that a rigorous simulation of the fiber suspension flow problem should include the effect of the flow on the fiber orientation as well as the fiber orientation influence on the flow rheology, in which the “fully-coupled” calculations require considerable computational expenses. To save computational time, these models are typically decoupled, such as in injection and compression modelling simulations, by using the flow kinematics to predict the fiber orientation.

To save computational time, these models are typically decoupled, such as in injection and compression modelling simulations, by using the flow kinematics to predict the fiber orientation. The decoupled formulation has been shown (c.f.1082Tucker et al. [6]) to work well in shear dominated flows such as those that occur in injection molding [7] and compression molding [8].Fiber orientation investigations involved in Additive Manufacturing applications are still quite limited. Nixon, et al. [9] numerically evaluated the fiber orientation conditions for three FDM nozzle geometries (convergent, straight and divergent) using the Moldflow (Moldflow Corporation, Framingham, MA) and the Folgar-Tucker diffusion model. The computed results showed that the convergent geometry achieved the highest fiber alignment and the divergent geometry resulted in the lowest.

The computed results showed that the convergent geometry achieved the highest fiber alignment and the divergent geometry resulted in the lowest. Additionally, at the exit of the straight nozzle as well as the convergent nozzle, a higher alignment is found around the core area than the boundary area which was not seen in the experimental result reported by Kunc [10]. Heller, et al. [11] predicted the fiber orientation tensor of a conventional FDM nozzle and the extruded filament where die swell was evaluated numerically in the solution process. The COMSOL Multiphysics (Comsol, Inc., Burlington, MA) software was used to calculate the velocity and the velocity gradient field of the flow domain assuming the polymer behaved as an isothermal Newtonian fluid in a creeping flow. Orientation tensors (see, e.g. Advani and Tucker [12]) were used to evaluate fiber orientation along streamlines within the flow domain from velocity and velocity gradient information.

Orientation tensors (see, e.g. Advani and Tucker [12]) were used to evaluate fiber orientation along streamlines within the flow domain from velocity and velocity gradient information. Their results show that fiber alignment reached its peak at the boundary of the nozzle and decayed towards to the center. Also, an extrudate swell ratio (c.f. Equation 7 below) of 13% was computed which is near the established accepted values for Newtonian fluids.Extrudate swell occurs in many extrusion-based polymer processes such as that observed for the Strangpresse (Strangpresse, Inc, Youngstown, OH) AM extruder appearing in Figure 2-a. Based on earlier results, it is expected that the sudden change of the flow boundary at the nozzle exit will result a notable variation in fiber orientation downstream. Molten polymers are well known to exhibit a non-Newtonian behavior which has been modeled successfully with viscoelastic fluid models that exhibit both viscous and elastic effects.

Molten polymers are well known to exhibit a non-Newtonian behavior which has been modeled successfully with viscoelastic fluid models that exhibit both viscous and elastic effects. The flow of a viscoelastic fluid is influenced by the first normal stress difference, which is considered to be a main contributor to die swell [13]. Keunings [14] and Huang, et al. [15] reviewed the preceding literature associated with the discussion of die swell. Luo and Tanner [16] exploited the Streamline Finite Element Method (SFEM) on solving the die swell problem. Luo and Mitsoulis [17] refined the SFEM by introducing a particle-tracking scheme along the streamlines and iterate the algorithm by Picard scheme. Beraudo, et al. [18] introduced a finite-element-based method to investigate the extrudate swell of LLDPE and LDPE melts using the Phan-Thien- Tanner (PTT) Model [22], a differential-type viscoelastic fluid model, based on geometries of a 2D slit die and a 2D axisymmetric capillary die.

Their results shown a good agreement with related experimental data but over-estimated the swell in high shear condition. Ganvir, et al. [21] proposed an Arbitrary Lagrangian Eulerian (ALE) algorithm for solving the die swell problem for both the steady state and transient problems. Their formulation behaved well at high shear rate but was unstable under low shear.This paper presents a numerical approach to evaluate fiber orientation in a large scale FFF extruder nozzle and the post-nozzle extrudate, including a prediction of the die swell at the nozzle exit. Our numerical model is created through the finite element suite ANSYS Polyflow (ANSYS, Inc., Canonsburg, PA). When modeling the flow, three Generalized Newtonian Fluid (GNF) models and a viscoelastic fluid model are employed in separate simulations. A decoupled1083formation is used to compute the fiber orientation within the polymer melt flow.

A decoupled1083formation is used to compute the fiber orientation within the polymer melt flow. The Advani- Tucker fiber orientation tensor [12] and the Isotropic Rotary Diffusion Model [19] are employed in the computation of the fiber orientation along streamlines within the flow domain. The Fast Exact Closure of Montgomery-Smith, et al. [20] is applied to address the closure problem in the fiber orientation simulations.Nozzle Flow ModellingThe simulations involved in this research are based on the large scale additive manufacturing Strangpresse Model-19 extruder appearing in Figure 1. Figure 1-a shows the extrusion of neat ABS polymer from the nozzle and important geometrical dimensions of the Strangpresse Model 19 large scale AM extruder nozzle appear in the Figure 1-b. Our extruder nozzle flow model is based on this design, and also includes a 20-mm strand of polymer extrudate beyond the nozzle exit to simulate the effect of die swell.

Our extruder nozzle flow model is based on this design, and also includes a 20-mm strand of polymer extrudate beyond the nozzle exit to simulate the effect of die swell. Due to the axisymmetry of the nozzle geometry and assumed flow, we consider only half of the flow domain using a 2-D axisymmetric model to reduce computational expense. Note that the axisymmetric assumption ignores any swirling motion of the flow that may result from the extruder screw, which is a topic of future research.The boundary conditions of the flow domain are shown in Figure 2-a, in which Q is the volumetric flow rate, F(cid:2929) is the tangential force, F(cid:2924) is the normal force, v(cid:2929) is the tangential velocity, v(cid:2924) is the normal velocity, (cid:1822) is the velocity vector at the free surface, and (cid:1814) is normal vector to the free surface [13]. The finite element domain is discretized into 20,382 nodes and 19,866 elements using the 4-node quadrilateral element.

The finite element domain is discretized into 20,382 nodes and 19,866 elements using the 4-node quadrilateral element. The mesh in the convergence zone and the nozzle exit is shown in Figure 2-b. Additionally, the polymer melt flow is under the isothermal assumption for simplicity.(a) (b)Figure 1. (a) Neat ABS exhibit obvious die swell as is extruded throught the Model-19 extruder nozzle. (b) Essentail geometric dimensions of the nozzle.1084(a) (b)Figure 2. (a) Essentail boundary conditions of the flow domain (b) Mesh of convergent zone as well as the nozzle exit.Material RheologyGanvir, et al. [21] simulated die swell of the LDPE and LLDPE in a capillary die using the Phan-Thien-Tanner (PTT) rheology model [22] for the molten polymer. The numerical prediction gave a good agreement with previous experimental and numerical results [21].

The numerical prediction gave a good agreement with previous experimental and numerical results [21]. The PTT model is assumed to be one of the most realistic rheology model in the family of the Differential Viscoelastic Fluid (DVF) models [13]. A comparable study using ABS has yet to be performed, so we have chosen to use LLDPE properties from Ganvir, et al. [21] in this study to validate our PTT flow modeling approach given that data exist in the literature for this comparison. However, we do realize that LLDPE is not common in AM, and our future work will focus on ABS and other AM materials as we develop a database of the required viscoelastic material properties.The shear viscosity of a polymer melt is well known to be shear rate dependent and commonly exhibit shear thinning behavior. This is particularly true for LLDPE [21]. To capture the shear thinning phenomenon in our nozzle flow model, we first consider various Generalized Newtonian Fluid (GNF) models.

This is particularly true for LLDPE [21]. To capture the shear thinning phenomenon in our nozzle flow model, we first consider various Generalized Newtonian Fluid (GNF) models. The Power law model is given as [23]η(cid:4666)γ(cid:4662) (cid:4667) (cid:3404) K(cid:4666)λγ(cid:4662) (cid:4667)(cid:2924)(cid:2879)(cid:2869)(Equation 1)1085where K is the consistency factor, λ is the natural time, and n is the power-law index. The Bird- Carreau law captures the shear thinning behavior of the Power law form, and also includes a Newtonian plateau at low shear rates.

Note that the Cross law also includes a Newtonian plateau at low shear rates. The shear viscosity data obtained from Ganvir, et al. [21] are fitted with each of these GNF models as is shown in Figure 3.Figure 3. Curve fitting the shear rate dependent viscosity with different GNF Laws.In addition to the GNF models above, we also consider the PTT viscoelastic rheology model in our study. Ganvir, et al. [21] represented LLDPE with a 7-mode PTT model, in which the essential parameters of the model are given in Table 1 [21].

The tensor D (cid:3404) (cid:2869) (cid:4670)(cid:4666)(cid:1487)(cid:1822)(cid:4667) (cid:3397) (cid:4666)(cid:1487)(cid:1822)(cid:4667)(cid:2904)(cid:4671) is the strain rate tensor, where (cid:1822) is the velocity vector, (cid:2019) is the mode relaxation (cid:2870) time, (cid:2015) is the mode viscosity, (cid:2022) controls the shear viscosity behavior and (cid:2013) controls the elongational behavior.Moreover, we plot the storage shear modulus, loss shear modulus and the steady state shear viscosity in Figure 4 based on the parameters given in Table 1. As is seen, the PTT model also presents the shear thinning behavior that is given in the GNF models above.From Equations 2 to 6, it can be seen that the Power Law, Bird-Carreau law, Cross law as well as the PTT model include inherent non-linearity in their constitutive equations.

We expect that polymer melt flow results using GNF and PTT rheology models may vary considerably from those obtained when assuming a Newtonian fluid model. Furthermore, simulations using non- Newtonian models are nonlinear which would yield flow-rate dependent results. In the study below, we use a volumetric flow rate Q =10 (cid:1865)(cid:1865)(cid:2871)/(cid:1871) with the fully developed velocity profile inlet condition imposed as is shown in Figure 5. Other flow rates are expected to yield different fiber orientation results when using non-Newtonian fluid models.Table 1.

In detail, the Power Law [23] gives a 1.8% swell ratio, Bird-Carreau Law and Cross Law [23] give similar results, which are about 8.8%. Specifically, the die swell predicted by the Cross law is 30.5% less than using the Newtonian model.Extrudate Swell Prediction using PTT ModelMelt flow simulations were also performed for the large scale FFF nozzle flow using the PTT viscoelastic fluid model. The ANSYS-Polyflow PTT model was employed using the LLDPE parameters appearing in Table 1 above. Before simulating the FFF nozzle with the PTT fluid, we first considered the 2D axisymmetric capillary flow domain shown in Figure 6-a taken from Ganvir, et al. [21] to validate our PTT flow simulation process. Note that the lower 14.2 mm of the model is used to capture the die swell. In this model, the die length to diameter ratio ((cid:1838)/(cid:1830)) is 19.2 and the volumetric flow rate is about 10 (cid:1865)(cid:1865)(cid:2871)/(cid:1871).

In this model, the die length to diameter ratio ((cid:1838)/(cid:1830)) is 19.2 and the volumetric flow rate is about 10 (cid:1865)(cid:1865)(cid:2871)/(cid:1871). It can be seen that the predicted free surface shown in Figure 6-b has a good agreement with the published literature.Once our modeling approach was validated with literature data above, we applied the PTT rheology model to the FFF nozzle flow domain shown in Figure 2-a to obtain the velocity field within the melt flow and the predicted die swell. From the comparison shown in Figure 7. Computed results show that the PTT model yields a higher die swell than the GNF models due to the elastic effect that is included in the PTT model. Moreover, the swell profile of the PTT model does not reach the steady state within the first 20 mm of the free extrudate while the results of Newtonian Fluid model and GNF models reach steady state at roughly the first 2 mm.

Note, the PTT model’s 22.4% swell ratio is computed based on the swelled diameter at the 20 mm position. In particular, the PTT model’s predicted result is 72.2% higher than that of the Newtonian Fluid model, and the increase between results of the PTT model and the Cross law is even over 160%.The above analysis indicates that the elastic effects of the viscoelastic fluid flow significantly enhance the extrudate swell phenomenon even with considering the decreasing effect of the shear thinning behavior.1089(a) (b) Figures 6. (a) Representation of the capillary die used in the validation study (b) Extrudate swell predictions for the die flow domain.Figure 7. Extrudate swell predictions of the PTT model, NF model, and GNF models.Fiber Orientation Prediction using Decoupled ModelThe flow-induced fiber orientation pattern is determined using the decoupled fiber orientation tensor formulation [11].

In addition, (cid:2019) is a constant that depends on the fiber aspect ratio, (cid:1829)(cid:3010) is the interaction coefficient that models the effect of fiber-fiber interaction on the orientation distribution. The scalar magnitude of the rate of deformation tensor D is (cid:2011)(cid:4662) and (cid:2414) is the 4th order fiber orientation tensor, which is typically solved by a closure approximation [3]. In this study, Fast Exact Closure [20] is used to solve for (cid:2414). It is important to note that the orientation tensor approach does not track each individual fiber, but instead provides an indication of the degree of orientation through the components of A appearing in Equation 9. As shown in Figure 8, values of a diagonal component of A having a value of one represents full alignment in a given direction (shown as (cid:1876)(cid:2870) in Figure 8-a), and values of 1/3 on all diagonal components represent the case of a uniformly random orientations (cf. Figure 8-b).

Figure 8-b). Recall, in Figure 2, we define the extrusion direction of the melt is along (cid:1876)(cid:2870) direction, thus, the (cid:1827)(cid:2870)(cid:2870) component is of particular importance since it indicates how well the fibers align along the principal extrusion direction.The fiber orientation tensor solution of A22 shown in Figure 10 is based on the flow kinematics solved using the PTT rheology model, in which the streamlines numbers in the legend corresponds to the highlighted streamlines in Figure 9. In our solution, we assume the initial fiber orientation state at the nozzle entrance is uniformly random, and we note that other initial conditions on A (at the inlet) may result in different fiber orientation variation within the nozzle geometry. In addition, we assume (cid:2019) = 0.9802 and (cid:1829)(cid:3010) = 0.0075, as in Heller, et al.

In addition, we assume (cid:2019) = 0.9802 and (cid:1829)(cid:3010) = 0.0075, as in Heller, et al. [11].A relative comparison is given in Figure 11, in which the steady state values of the (cid:1827)(cid:2870)(cid:2870) component solved by flow kinematics using the Newtonian fluid model, GNF models and the PTT model. Among the comparison, it shown that the PTT model kinematics yields lowest steady state principal fiber orientation alignment, on average. The variation of the principal fiber orientation alignment decrease initially and recovers gradually as moving from the boundary toward to the center of the nozzle. Furthermore, we compute the absolute relative difference (∆,1091cf.

(cid:1827)(cid:2870)(cid:2870) component of fiber orientation solution computed using the flow kinematics solved by the PTT model.Figure 11. Comparison between the steady state (cid:1827)(cid:2870)(cid:2870) component solved by different flow kinematic results.1093ConclusionA few conclusions have been reached based on the simulations presented in this paper. The shear thinning behavior of polymer melts reduces die swell effect. In particular, in the case of LLDPE [21], the reduction of the swell ratio is 30.5% comparing GNF Cross law to the Newtonian fluid model. The viscoelastic rheology of polymer melts greatly enhances the die swell.

The viscoelastic rheology of polymer melts greatly enhances the die swell. Again, by using LLDPE [21], we found an increase of 72.2% in terms of the swelled ratio, comparing the PTT model results to the Newtonian fluid results.Based on the fiber orientation prediction of the decoupled model, it is observed that the highest principal alignment state occurs at the streamline closest to the boundary, and shows a concave down variation trend as moving from boundary to the center line. Furthermore, the PTT model yields lower principal alignment results than those from GNF models. In detail, the prediction of the PTT model is 7% lower comparing to Newtonian model’s; 13% lower comparing to Bird-Carreau’s or Cross law’s; and 26% lower comparing to the Power law, on average.Future WorkFuture work will include simulations with materials and flow rates that are typical of large scale polymer deposition additive manufacturing.

This includes ABS and carbon fiber filled ABS materials, and flow rates in the 10 to 15 lbs per hour range. We also plan to develop a simulation method that employs the ANSYS Polyflow simplified viscoelastic rheology model to reduce computation time. Finally, a couple fiber orientation and flow velocity analysis is needed to more accurately account for influence of the suspended fibers on melt viscosity.AcknowledgementsThe authors would like to thank the Oak Ridge National Lab (RAMP-UP 40001455134) as well as Baylor University for funding this research project. Also, we would like to thank the Strangpresse corporation for donating their Model-19 extruder to our study.References[1] Brenken, B., Favaloro, A. Barocio, E. Denardo, N. Kunc, V., Pipes, R.B., Fused Deposition Modeling of Fiber-Reinforced Thermoplastic Polymers: Past Progress and Future Needs. In Proceedings of the American Society for Composites: Thirty-First Technical Conference 2016.

In Proceedings of the American Society for Composites: Thirty-First Technical Conference 2016. [2] Love L.J., Utility of big area additive manufacturing (BAAM) for the rapid manufacture of customized electric vehicles. Oak Ridge National Laboratory (ORNL); Manufacturing Demonstration Facility (MDF); 2015 Aug 1. [3] Verweyst B.E., Tucker C.L., Fiber suspensions in complex geometries: flow/orientation coupling. The Canadian Journal of Chemical Engineering. 2002 Dec 1;80(6):1093-106. [4] Evans J.G., The effect of non-Newtonian properties of a suspension of rod-like particles on flow fields. Theoretical Rheology, Halstead Press, New York. 1975:224-32. [5] Lipscomb, G.G., Denn, M.M., Hur, D.U., Boger, D.V., The flow of fiber suspensions in complex geometries. Journal of Non-Newtonian Fluid Mechanics. 1988 Jan 1;26(3):297-325. [6] Tucker C.L., Flow regimes for fiber suspensions in narrow gaps. Journal of Non-Newtonian fluid mechanics.

This work describes the method of determining the effect of porosity that resembles 3D printed electronics. In addition to mechanical strength, determination of desirable limit of electrical insulation and heat dissipation will allow the additive manufacturing community to fabricate power electronics components with reduced cost and improved performance. For experimentation, three different sets of coupons were fabricated using Polycarbonate (PC) thermoplastic with embedded bare copper wire. Characterization included high electrical stresses and thermal testing to determine the effect of porosity on insulation and heat dissipation, respectively. During electrical characterization, higher wire density resulted in reduced breakdown strength.

During electrical characterization, higher wire density resulted in reduced breakdown strength. In thermal test, the comparisons between as fabricated and heat-treated specimen showed that heat dissipation increased by an average of 30 % to 40 %.IntroductionIn the polymer material extrusion additive manufacturing (AM) community, there is an immense interest in the fabrication of multifunctional parts by incorporating two or more manufacturing processes. An example of these processes is machining which is used to improve the surface finish of fused deposition modeling (FDM) built substrate, enhancing the quality of the part. Both operations could be done as post-processing. Printing and interrupting the build is a luxury one has with an FDM system as it allows the manufacturer to introduce new component as well as conventional machining if necessary.

Printing and interrupting the build is a luxury one has with an FDM system as it allows the manufacturer to introduce new component as well as conventional machining if necessary. The pausing advantage can be taken in a continuous AM process by incorporating a second manufacturing platform that will allow the further customization on fabricated parts. For this type of component integration platform, the Multi3D foundry system is a prime example [1]. The foundry system includes the introduction of an electrical component, a pick and place, machining, and embedding circuitry elements within the printed parts.

The foundry system includes the introduction of an electrical component, a pick and place, machining, and embedding circuitry elements within the printed parts. For example, wire embedding on the fused deposition modeling (FDM) part is one the processes of increasing the part functionality such as electrical, mechanical, and thermal.While the improvement of mechanical properties of FDM built parts is taking a continuous effort in designing new materials, and the development of process parameters, porosity remains as an inherent challenge for many applications. Several factors such as resolution, the temperature of the printing environment, and thermal gradient within the rasters are responsible for the formation of voids. Although tissue engineering application was the primary focus, authors in [2] showed that the raster gap size has a dominant effect on the porosity, pore diameter, and the mechanical strength of the part.

For example, ABS specimen with 0.1 mm raster gap had a compressive strength of 14 MPa and doubling the raster gap to 0.2 mm reduced the strength to 8 MPa. On the contrary, 30 % porous (vol.) ABS specimen had compressive strength 17 MPa, while the specimen with 60% porosity had a value of 8 MPa. In a separate design experiment, it showed that every 20 % (vol.) increasing of porosity within ABS specimen was reducing the mechanical strength by 5 MPa [3]. Therefore, it is obvious that changing printing parameters is not enough to avoid the porosity within parts; improvement of the mechanical properties requires further post-processing procedures.340Figure 1. (a) CAD for fabricated coupon and (b) fabrication of wire embedded coupon using thermal wire embedding toolKnowing the presence of voids in printed part, tremendous effort is being made to increase the application- based AM, specifically for 3D electronics fabrication, however, the effect of porosity remains unexplored.

One of the major challenges of porosity on electrical parts is the failure of breakdown strength. Breakdown strength, also known as high electrical potential (hipot), is the ability to withstand sudden rise of voltage of an insulating material. For example, in medium voltage application (300~1000V), breakdown strength is an essential parameter to ensure the reliability of the manufactured part. Another challenge of the porosity is the decrease in thermal conductivity of a printed part. Within the part, air has a conductivity value of 0.02 W/m/K, and plastic varies from 0.18~0.22 W/m/K. Using Voigt’s rule, the increase of the porosity results in a lesser amount of combined thermal conductivity in the part.

Using Voigt’s rule, the increase of the porosity results in a lesser amount of combined thermal conductivity in the part. Knowledge of heat dissipation on FDM parts will explore a new dimension in 3D electronics fabrication with advanced and efficient thermal management.Although the mechanical properties are heavily demonstrated by myriads of article in past, this research was aimed to investigate the effect of porosity on electrical and thermal properties of an FDM built parts. Particularly in 3D electronics fabrication and characterization, the necessity of knowing the effect porosity has on breakdown strength and heat dissipation is undoubtedly essential to explore new product development and production. Additionally, the findings of this research would help to improve the thermal management plan, electrical safety, and operational reliability.

Additionally, the findings of this research would help to improve the thermal management plan, electrical safety, and operational reliability. To accomplish these specific set of goals, coupons were fabricated using polycarbonate (PC) with material extrusion AM and embedding bare copper wires followed by characterizations: high potential insulation testing and thermal testing including heat treated and untreated test specimens. Lastly, it was anticipated that findings from these experiments can aid in designing future hybrid additively manufactured power electronics (such as stators, inductors for a wireless charger, or embedded transformers).FabricationTo investigate the effect of porosity in electrical and thermal performance of FDM parts, three sample sets, each containing fifteen specimens, of wire embedded coupons were fabricated. In each sample set, ten specimens were chosen and had either one, two, or three wires embedded within each specimen.

In each sample set, ten specimens were chosen and had either one, two, or three wires embedded within each specimen. Of the ten specimens within each group, five were tested in an as-fabricated condition while the other five specimens were heat treated before testing. The heat treatment consisted of placing the specimens in an oven for two hours. The oven temperature was set at 165 ºC, which was above the glass transition temperature of PC (161 ºC), to promote polymer chain mobility and reduction of voids while limiting temperature induced dimensional changes, which341were more pronounced above 165 ºC.

The 3D printing was accomplished by using a Multi3D foundry system which can be read in details at [1].In this research, a FDM Fortus 400Mc machine (Stratasys, Eden Prarie, MN, USA) in Multi3D foundry system was used to deposit polycarbonate (PC) (Stratasys, Eden Prarie, MN, USA) and embed a series of 26 AWG (Ø = 0.405 mm) bare copper wires (Arcor Electronics, Niles, IL, USA) in an automated fashion. The embedding of bare copper wire within the PC was accomplished in following steps (i) dispensing PC material to produce a substrate, (ii) transferring the workpiece via the six-axis robot to the CNC machine where wire embedded using a cartridge heating wire embedding tool, and (iii) transferring the workpiece back to the FDM printer to dispense PC onto the substrate. Figure 1 (a) is a computer-aided design (CAD) representations for each of the three sample groups. The 3D printing layer thickness was set at 0.254 mm (0.010 inch).

Figure 1 (a) is a computer-aided design (CAD) representations for each of the three sample groups. The 3D printing layer thickness was set at 0.254 mm (0.010 inch). The coupon was fabricated by stacking six PC layers before embedding the copper wire, which was embedded within the sixth PC layer and occupied approximately half the thickness of the fifth PC layer. In Error! Reference source not found. (b), wire embedding on PC substrate at CNC router is shown. A build interruption in FDM printing was made to transfer the build platform to CNC router and embedded the wire.Experimental MethodologyThree characterization techniques were used to determine the effect of porosity including (i) hipot testingusing AC or DC, (ii) thermal testing, and (iii) cross-sectional micrograph analysis.Hipot testingAccording to IEEE Std. 1995 1995 [6], both AC and DC hipot testing was done in this work.

1995 1995 [6], both AC and DC hipot testing was done in this work. AC hipottesting voltage was chosen by the following equation:𝑉𝑇𝑒𝑠𝑡 = (2 × 𝑉𝑂𝑝𝑒𝑟𝑎𝑡𝑖𝑜𝑛𝑎𝑙 + 1000)In the case of DC hipot testing, the test voltage was chosen using the following equation:{2} A HiPot-3870 Tester with device under testing (DUT) enclosure (Associated Research Inc., Lake Forest, IL, USA) was used to conduct the experiment. The conductive copper wire was connected to the high voltage electrode while the insulating polycarbonate was connected to the low voltage electrode of the hipot tester where “Test” and “Operational” are the expected and end-use operating voltage, respectively. During AC hipot test, the considered 𝑉𝑜𝑝𝑒𝑟𝑎𝑡𝑖𝑜𝑛𝑎𝑙was 2,000 V (ramp 0.1s) and dwell time of the test was 10 s. The DC test was performed similarly to the AC test by considering 𝑉𝑜𝑝𝑒𝑟𝑎𝑡𝑖𝑜𝑛𝑎𝑙 as 970 V, which resulted in a 𝑉𝑇𝑒𝑠𝑡 of 5 kV.

The DC test was performed similarly to the AC test by considering 𝑉𝑜𝑝𝑒𝑟𝑎𝑡𝑖𝑜𝑛𝑎𝑙 as 970 V, which resulted in a 𝑉𝑇𝑒𝑠𝑡 of 5 kV. In the DC test, the ramp was 5 s and total dwell time was 1 minute.𝑉𝑇𝑒𝑠𝑡 = (2 × 𝑉𝑂𝑝𝑒𝑟𝑎𝑡𝑖𝑜𝑛𝑎𝑙 + 1000) × 1.7Thermal testingThe steady state surface temperature of the printed thermoplastic was determined by simulating the Joule heating effect within the embedded wire using ANSYS steady state thermal module (ANSYS, Inc., Canonsburg, PA).

Boundary conditions in the modeling included using a convective heat transfer coefficient (h) of 12, 15, and 15.27 W/m2/°C for all six sides of the one-, two-, and three-wire thermoplastic coupon respectively, an effective conductive heat transfer coefficient (K) of 0.182 W/m/K (using Voigt’s model)Thermal testing was performed using the Joule heating principle to generate heat within the embedded copper wire (26 AWG or Ø = 0.405 mm) by supplying current (5 A) from a DC power supply (BK PRECISION 9115, B&K Precision Corporation, Yorba Linda, CA, USA). Two K-type thermocouples were placed on the top and bottom surface of the specimen to collect the surface temperature. Data acquisition unit NI – cDAQ 9426 (National Instruments, Austin, TX) was used to measure temperature. Data were acquired every second for a total duration of two hours.

Data acquisition unit NI – cDAQ 9426 (National Instruments, Austin, TX) was used to measure temperature. Data were acquired every second for a total duration of two hours. For each test specimen, three separate tests were performed at different times to verify the repeatability of the experimental data.342{1}Figure 2. AC hipot testing before and after heat treatment of wire embedded couponsResults and discussionElectrical characterizationAs it mentioned earlier that electrical breakdown strength was measured in both AC and DC test. AC hipot testing of the coupons containing embedded wires was designed to test the ability to withstand low and medium voltages. To keep in mind that, it was not the goal to determine the dielectric strength of the PC material, rather it was to observe the effect of voids in insulating materials. Figure 2 shows the average breakdown strength (AC hipot testing) from five measurements for each group before and after heat treatment.

Figure 2 shows the average breakdown strength (AC hipot testing) from five measurements for each group before and after heat treatment. The measured breakdown strength ranged for one-, two-, and three-wire untreated specimens (with constant thickness of PC at 2.5 mm) were 4 - 4.8 kV, 1.98 - 2.6 kV, and 2 - 2.8 kV, respectively. It was expected that breakdown strength of coupons would decrease with increasing number of wires, which was found in one-wire and two-wire specimens, however, the comparisons of the two-wire and three-wire specimens did not agree with the expectation. It was assumed that supper position of electric field is responsible for lowering the breakdown strength.Breakdown strength of heat-treated specimen was increased as shown in Figure 2. Among the 5 tests, four of the one-wire breakdown strength was above 5 kV. In the case of the two- and three-wire specimen, average breakdown strength was 4 – 4.03 and 4.15 – 4.3 kV, respectively.

Among the 5 tests, four of the one-wire breakdown strength was above 5 kV. In the case of the two- and three-wire specimen, average breakdown strength was 4 – 4.03 and 4.15 – 4.3 kV, respectively. During DC hipot testing, all the specimens from the three sample groups passed (i.e., did not breakdown) when subjected to 5kV in both the heat-treated and untreated condition.Thermal characterizationThe steady-state substrate surface temperature of one-wire specimens was simulated 51 °C Figure 3(a). The empirical surface temperature for the same specimens without heat treatment was measured as 38°C. After the heat treatment a steady state surface temperature of 51 °C was observed as shown in Figure 3(b). While the heat-treated surface temperature was in close agreement with the simulated temperatures, the non-heat-treated surfaces was substantially lower.

While the heat-treated surface temperature was in close agreement with the simulated temperatures, the non-heat-treated surfaces was substantially lower. The mismatch is attributed to the inherent porosity of printed parts, which influences the conductive heat transfer coefficient.In the case of two-wire and three-wire specimens shown in Figure 3(c) to (f) the simulated and empirical substrate surface temperatures had close agreements. Since the number of wire increased, relatively higher temperature was measured on the PC surface. Besides the PC surface temperature, a sound agreement of analytical 343Figure 3. Temperature contours and experimental temperature plot of wire embedded coupons.and simulated temperature of the embedded conductor surface was found. All sets of thermal testing results showed agreement of temperature difference between heat-treated and untreated specimens as well as with the analytical temperature.

All sets of thermal testing results showed agreement of temperature difference between heat-treated and untreated specimens as well as with the analytical temperature. Now it is obvious that porosity reduction increases the heat conductivity because air has a lower thermal conductivity (0.02 W/m⋅K) which was filled up by the expansion of each rod of PC during heat treatment.Micrograph analysisTo further confirm the reduction porosity and subsequent effect on breakdown strength and heat dissipation of heat-treated specimen, micragraph analysis was performed. Scanning Electron Microscopy (SEM) images of heat-treated and untreated 3D printed specimens revealed reduction of porosity. For example, thermal testing of untreated specimen confirmed that the presence of porosity reduces the heat dissipation whereas the heat-treated specimen having less porosity also had higher heat dissipation.

The higher heat conduction was the result of reduced voids by the expansion of PC rod and ultimately the bulk of PC. In Figure 4 the presence of porosity is in heat treated and untreated specimen is shown. Therefore, the main finding was the reduced porosity led not only the higher heat dissipation but also increased the breakdown strength.344Figure 4. SEM image of cross sectioned specimen before heat treatment (a) one-wire, (b) two-wire, (c) three-wire and after heat treatment (d) one wire, (e) two-wire, and (f) three-wire.ConclusionThis research paper described the effect of porosity in electrical and thermal performance of material extrusion 3D printed parts containing embedded solid conductor. Test coupons containing PC substrate material and embedded bare copper wire were fabricated using an automated hybrid additive manufacturing process via the Multi3D system.

Test coupons containing PC substrate material and embedded bare copper wire were fabricated using an automated hybrid additive manufacturing process via the Multi3D system. Through experimental testing of wire embedded coupons that has voids, the ability to withstand at high electrical stress was determined. Thermal testing was conducted to confirm the hypothesis of improving the porosity by heat treatment. Following three conclusions were made from the present studies to strongly support the claim of effect of porosity on electrical insulation and heat dissipation of 3D printed parts:1. Heat treatment of wire embedded specimen increased the breakdown strength of each group. For example, two-wire untreated specimen had average AC breakdown strength 2.4 - 2.8 kV while after heat treatment it was 4 – 4.2 kV. Therefore, the effect of porosity was found in breakdown strength by conducting tests before and after the reduction of voids in printed parts.2.

Therefore, the effect of porosity was found in breakdown strength by conducting tests before and after the reduction of voids in printed parts.2. Untreated specimen had inherent porosity that led the lower heat dissipation. Heat treated specimen had higher heat dissipation. This was quantified in each case such as the three-wire coupons were substrate temperature increased from 56°C to 71°C after reducing porosity by heat treating. Additionally, simulated and empirical data had good agreement in terms of surface temperatures of each group. 345AcknowledgementThe authors would like to express appreciation for the National Aeronautics and Space Administration (NASA) who partially funded this work under Grant NNC17CA02C. The fabrication and characterization presented here was conducted at The University of Texas at El Paso (UTEP) within the W.M.

The fabrication and characterization presented here was conducted at The University of Texas at El Paso (UTEP) within the W.M. Keck Center for 3D Innovation (Keck Center) using equipment developed via a previously funded effort through the National Center for Defense Manufacturing and Machining under the America Makes Program entitled ‘3D Printing Multifunctionality: Additive Manufacturing for Aerospace Applications’ (Project #4030) and based on research sponsored by Air Force Research Laboratory, under agreement number FA8650-12-2-7230. The authors acknowledge the contribution of Carlos Acosta, Lluvia Herrera, Sol Barraza, and Leonardo Gutierrez for their assistance and support during experimentation.The views and conclusions contained herein are those of the authors and should not be interpreted as necessarily representing the official policies or endorsements, either expressed or implied, of Air Force Research Laboratory or the U.S. Government. The U.S.

FECT OF POST-PROCESSING ON THE MICROSTRUCTURE AND MECHANICAL PROPERTIES OF ULTRA-LOW CARBON STEEL FABRICATED BY SELECTIVE LASER MELTINGB. Almangour, J. M.YangDepartment of Materials Science and Engineering, University of California Los Angeles, Los Angeles, CA 90095, USAAbstractIn this study, the effects of heat treatments and hot-isotactic pressing (HIP) on the microstructure and mechanical properties of ultra-low carbon steel produced using selective laser melting (SLM) were investigated. Powder and prototypes characterizations including XRD phase analysis, microstructural observations, and hardness were performed. It was found that heat treatments at 1000 ºC and HIP process improved inter-particle bonding very slightly.

It was found that heat treatments at 1000 ºC and HIP process improved inter-particle bonding very slightly. Significant increases in the grain size were observed for the annealed specimens at 600ºC and above as well as after HIP due to recrystallization and further grain growth, which coincide with the drop in hardness.Keywords: Additive manufacturing; selective laser melting; porosity; heat treatment; HIPIntroductionAdditive Manufacturing (AM) is a group of advanced technologies used to fabricate engineering prototypes of parts through layers material addition based on a digital 3D design data of the item. [1]. The starting material feedstock are shaped and consolidated to build up an arbitrary configuration, given that a computer model of the part has been generated on a computer-aided design (CAD) system [2].

The current trend exhibited in the field of AM is producing of complex and flexible functional components so that they may meet the demands, standards and qualities required by several applications including automotive, aerospace, biomedicine, and the defense [3]. The unique advantage of AM is the possibility to build geometries that are very difficult to fabricate using traditional manufacturing techniques due to its additive approach; i.e. design freedom [4, 5]. This makes it less costly to print because it can be made faster. It is also more flexible because it can be changed in any format in reference to the original CAD file [3, 4]. AM depends on market demand to produce components that are relatively fully dense with mechanical properties like those of the bulk materials [6]. Selective Laser Melting (SLM), in which it is part of AM technologies, powders are heated to just below their melting point to facilitate bonding and reduce distortion of the finished product [7, 8].

Selective Laser Melting (SLM), in which it is part of AM technologies, powders are heated to just below their melting point to facilitate bonding and reduce distortion of the finished product [7, 8]. One characteristic with SLM is the refinement of the microstructure through increasing the rate of cooling rate, and increases oxidation through restricting removal of oxygen [9-11], which results in higher strengthening.Numerous scientific researches with SLM of iron-based alloys show several challenges. The finishing of the SLM of iron-based materials is said to be rough [12]. Indeed, there is no uniform distribution of its particles and components causing a weak bonding within the matrix [13]. In general, the densification rate of the SLM of iron parts increased as the density of the applied laser energy increased.

In general, the densification rate of the SLM of iron parts increased as the density of the applied laser energy increased. However, increased reaction temperature (because of higher laserpower) and the inert atmosphere culminate to decomposition of Fe2O3 into FeO and Fe and oxygen that escapes from the specimen [6]. On this basis, higher laser power (causing higher temperatures/heating) contributes to removal of oxygen through intensifying the melting, decomposition, and evaporation. During the processing of SLM of steel, the microstructure suffers from several defects mainly porosity and shrinkage, which are very undesirable solidification defects [7, 14-18]. Besides, the loss of ductility and development of residual stresses may not be desirable for many applications.Post-processing treatments offer significant improvement to the density of the objects and hence to the mechanical properties.

It improves distribution and particles consolidation resulting in a strong particle-particle bonding [11, 19-21]. Heat treatments and Hot-Isotactic Pressing (HIP) therefore could help in eliminating some of the side effects of SLM, discussed above. It is an attempt to increase the strength and ductility by properly varying their microstructure. Limited attention has been made on studying the influence of heat treatments and HIP on the microstructure and mechanical properties of SLM materials.

Limited attention has been made on studying the influence of heat treatments and HIP on the microstructure and mechanical properties of SLM materials. Therefore, in this study, the effect of annealing at different temperatures as well as HIP treatment on the microstructure, phase distribution, and the mechanical properties using compression tests and hardness tests are investigated; for the purpose of comparison a conventional sample are tested using the same characterization tools.Experimental proceduresFeedstock powder, and SLM processesThe starting powder is low carbon steel powder (Hoeganes Company), which displays a general irregular shape (Fig. 1a). The composition of the feedstock is an iron matrix with 0.07 wt. % C (Table 1). The sizes of grains in powder particles were around 2µm, observed from the single etched particle (Fig. 1b). From the ground and polished particles, we can confirm that the particles are almost dense (Fig. 1c).

1b). From the ground and polished particles, we can confirm that the particles are almost dense (Fig. 1c). The particle size distribution of the powder shows a mean value of about 20 µm (Fig. 2).Processing of samples by SLM was carried out using an SLM 250 HL machine (SLM solutions) equipped with an Yb-YAG laser. The working chamber provides a closed environment filled with an argon-gas in order to prevent the oxidation of steel during the fabrication process. A steel plate was used as the building platform. Before being installed, this platform was grit- blasted with alumina. Using several trials, small steel cubes measuring 5mm x 5mm x 5mm dimensions were finally fabricated using the processing parameters shown in Table 2.Heat treatments and hot-isotactic pressing (HIP) processIsothermal heat treatments of 1 h were performed in air using an electric resistance furnace and subsequently air cooled.

The numbers of investigated heat treatment temperatures are: 400, 800, 600, and 1000 ºC for microstructure and mechanical properties characterizations. The HIP processing takes place at 1162 ºC with a ±3 ºC tolerance.

The HIP processing takes place at 1162 ºC with a ±3 ºC tolerance. The operating pressure is 14,750 Psi with a ±250 Psi tolerance while the treatment time is 3 hrs ±15 minutes.a100 µmb30 µm50 µmc Figure 1: (a) SEM micrograph of ultra-low carbon steel powder; (b) single etched particle; (c) polished particles.Figure 2: ultra-low carbon steel particle size distributionTable 1: Ultra-low carbon steel compositionC 0.07 maxMn 0.04 maxO 0.26 maxS 0.009 maxSi 0.06 maxCr 0.09Ni 0.01Fe BalanceTable 2: SLM processing parametersPower for area volume175 WAtmosphere Scanning speed for area volume 688 mm/sec 30µmLayer thicknessArgonMicrostructural characterizationsSamples were mechanically grounded and polished with progressively finer grit papers using the standard metallographic techniques and etched in 3% Nital reagent.

The surface and the cross-sectional microstructures of the steel cubes were observed using a Nova 230 Nano scanning electron microscope to characterize the powder and fabricated samples. XRD measurements were conducted on a Bruker- binary V1 instrument, operating with a cobalt anticathode at 35kV and 40 mA; the scan step time is 25.6 min in the angular range (2θ) 30-90º.Mechanical testingVickers indentations was performed (Leco, LM800AT) at a small load of 10 g at 15 random particle locations. The hardness of the as-built, annealed, and HIP samples were also measured using a Vickers microhardness tester at a load of 100g for a minimum of 15 measurements taken at random locations on the polished surface.

Nano hardness on the polished powder, as well as as-built, annealed, and aged samples were also performed using an MTS nano indentation tester at room temperature.ResultsMicrostructural characterizationsFigure 3a shows the microstructure of the as-built sample, showing some cracking, which suggest that high thermal stresses and thermal deformation were induced during the process. This means that the thermal stresses created during the high heat energy input resulting from low scan speeds have more effect than the C-content on inducing cracks. Besides, carbon may results in some brittleness (0.07% compared relatively to 0.03% found in the 316L stainless steel). In general, the energy available for consolidation is considered to be reduced by two effects. First, partial reflection from the previously scanned line, and secondly, the need to remelt previously proceed material. Only minor porosity was removed after implementation of HIP as shown in Fig. 3b (see dashed circle).

Only minor porosity was removed after implementation of HIP as shown in Fig. 3b (see dashed circle). HIP was performed at high temperature but still below the melting point, so there is no liquid phase formed during the HIP process. Also, HIP was performed under isotactic pressure, so there is a minimal effect of gravity in aiding the densification by flow of liquid in the direction of gravity (i.e. the thickness or build direction).a1 mmb1 mmFigure 3: Cross-sectional microstructure of ultra-low carbon steel: (a) as-built; (b) after HIP.The etched microstructures of as-built and after HIP treatment from top and side views are shown in Fig. 4. The as-built sample shows very fine grain size as a result of fast cooling. It was seen that grains are of different sizes, may be due the varying thermal heat fluxes during every scanning tracks. After HIP, uniform coarse equiaxed grains were developed and annealing twins throughout the structure (Fig. 4c, Fig.

After HIP, uniform coarse equiaxed grains were developed and annealing twins throughout the structure (Fig. 4c, Fig. 4d).At lower heat treatment temperatures of 400 ºC and 600 ºC, the structure remain unchanged relative to the as-built sample in Fig. 5b. A relatively uniform microstructure of equiaxed grains suggested that recrystallization was complete at 800 ºC with further grain coarsening obtained after annealing at 1000 ºC (Fig. 5d).a30 µmb10 µm100 µm100 µmdc Figure 4: Etched cross-sectional microstructure of ultra-low carbon steel: (a) as-built (top view); (b) as-built (side view); (c) after HIP (top side); (d) after HIP) side view).a10 µmb10 µm10 µm10 µmc Figure 5: Etched cross-sectional microstructure of ultra-low carbon steel: (a) heat treated 400 ºC; (b) heat treated 600 ºC; (c) heat treated 800 ºC; and (d) heat treated 1000 ºC.dFor comparison purposes, it is clearly shown that the cast sample in Fig.

6b displayed very coarse grains with a mean grain size of 100µm compared with the as-built sample with a mean grain size of 3µm. This is attributed to the fast cooling in the SLM process as indicated earlier.10 µm400 µma Figure 6: Etched cross-sectional microstructure of ultra-low carbon steel: (a) as-built fabricated by SLM; (b) as-cast (conventional).bXRD analysisFigure 7 shows the typical XRD histogram for the powder, as-built, and after HIP treatments. It can be seen that the characteristic BCC iron (α phase) crystalline structure peaks are present in the powder. The XRD pattern of the as-built samples exhibits a significant peak broadening due to the grain refinements as a result of fast cooling. In addition, same set of peaks are displayed in the XRD pattern of the as-built sample indicates that there was no phase transformation or oxidation during fabrication.

In addition, same set of peaks are displayed in the XRD pattern of the as-built sample indicates that there was no phase transformation or oxidation during fabrication. From the XRD results, no peaks responsible for the carbides can be found in the powder or after fabrication of cubes. There are two reasons why the XRD could not detect carbides. First, the carbon contents are very low and the amount of carbides formed is low. Secondly, due to the quick solidification nature of the SLM process, the dendrites and cells are small and the sizes of the black spots are extremely small. As such, carbide peaks may be too weak to be detected compared to the inherent background noises from the XRD. The XRD profiles after HIP treatment shows typical polycrystalline structure.Figure 8 shows the XRD pattern for the cube samples after heat treatments. Annealing at 400ºC or 600ºC is similar to the as-built with a slight increase in grain size based on “Williamson-Hall” analysis.

Annealing at 400ºC or 600ºC is similar to the as-built with a slight increase in grain size based on “Williamson-Hall” analysis. However, from annealing of 800ºC and 1000ºC, additional peaks were observed.It can be seen that the bases of the peaks in the SLM parts are relatively wider than the powder and cast sample (Fig. 9). This suggests that the crystal lattice structures of SLM parts are experiencing certain level of internal stresses that are thermally induced during the rapid solidification of SLM process as discussed earlier.

casting).Mechanical behaviorFigure 10, 11 demonstrates the effect of heat treatment temperatures and HIP treatments of the ultra-low carbon steel on microhardness and nanohardness respectively. The hardness of the as-built increased compared to those of the feedstock powders due to the powder melting followed by rapid solidification. Both scale hardness show stable hardness up to 600 °C, but the hardness drops significantly after annealing at 800 °C; it reduced even further after annealing at 1000 °C as a result of grain growth.

Both scale hardness show stable hardness up to 600 °C, but the hardness drops significantly after annealing at 800 °C; it reduced even further after annealing at 1000 °C as a result of grain growth. After the HIP treatments, the microhardness shows significant drop and similar value for both top and side views because of the homogenization.90) V H(s s e n d r a h o r c iM260 240 220 200 180 160 140 120 100 80 60 40 20 0Figure 10: Variation of microhardness as a function of heat treatment temperature and HIP treatment.3) a P G(2.5s s e n d r a h o n a N21.51Figure 11: Variation of nanohardness as a function of heat treatment temperature and HIP treatment.DiscussionIt was deduced, from various research literature, that there was a great challenge of producing carbon steel samples fabricated by SLM, with full density. The difficulty is the result of the special chemical properties of crucial elements within the steel composition.

The difficulty is the result of the special chemical properties of crucial elements within the steel composition. Iron, which serves as a matrix element of steel, is very reactive to oxygen. Therefore, a certain degree of oxidation is unavoidable under normal SLM conditions and powder handling [22]. More so, during laser processing, there is a strong likelihood of occurrence of the “balling phenomena”. This is the result of contamination layer of oxide present on the surfaces of steel melt. Although the carbon content in this study is very low, it is imperative to note that the carbon content of steel plays a critical role in determining processability of SLM. High carbon content also has adverse effect on SLM produced steel, since they portray a limited densification response. Wright et al. [23], found out that with increase in carbon contents, there is an increase in the thickness of the carbon layer accumulated on the melt surface.

Wright et al. [23], found out that with increase in carbon contents, there is an increase in the thickness of the carbon layer accumulated on the melt surface. Just like the oxide layer, the carbon layer has a critical influence mostly on diminishing wettability and causing the melt to spheroidise rather than flow across the underlying surface. In addition to this, the brittleness of SLM manufactured high carbon content steel is increased by the formation of complex interfacial carbides at grain boundaries [24]. Childs et al. [25] through their research, concluded that the dissolution of carbides is favored by increasing the heat flow in the treated powder. On the same breath, increasing heat flow homogenizes distribution of alloying elements. Thus, a thin powder layer thickness, which is less than 100mm, is more recommendable for laser melting. This is besides the optimization of laser types and parameters.

Thus, a thin powder layer thickness, which is less than 100mm, is more recommendable for laser melting. This is besides the optimization of laser types and parameters. Due to this reason, high volumetric energy density for both elemental homogeneity and powder consolidation is achieved [26-28].During the laser melting process, the metal powder absorbs the laser energy, raises its temperature, melt, and then solidifies when laser moves away, the whole process is less than one second. The heat also transfers away from the melt zone quickly. This causes the melt to cool down rapidly leading to short arrangement time and consequently porosity. In addition, because of the high cooling rate, the gas adsorbed on the surface of the powder, as well generated because of the reaction during the laser melting, have no time to escape from the molten pool and hence porosity was formed.When free interfaces exist in the system (e.g.

liquid-liquid, liquid-gas interfaces) the temperature and compositional gradient will impose a surface tension gradient along the interface, which exerts a shear stress on the fluid. This induces a flow toward regions with higher values of γ termed Marangoni convection [14]. It is observed that during scanning “balling” has occurred. This is caused by surface tension driven by Marangoni convection (i.e. melting instability). Because of the strongly concentrated energy source, a large temperature gradient is often induced on the surface. If the surface tension γ depends on temperature, then the temperature gradient induces a surface flow that can have a considerable influence on the melting pattern. If the surface tension increases with temperature (dγ/ dT > 0), the liquid moves toward the energy beam on the surface, inducing narrower and deeper pool.

If the surface tension increases with temperature (dγ/ dT > 0), the liquid moves toward the energy beam on the surface, inducing narrower and deeper pool. However, if the surface tension decreases with temperature (dγ/ dT < 0), the surface flow moves away from the energy beam, leading to a wider and shallower pool.ConclusionsUltra-low carbon steel powder has been successfully used in the SLM process, and the effects of heat treatments and hot-isotactic pressing (HIP) on the microstructure and mechanical properties were studied. Only minor porosity was eliminated after implementation of HIP and heat treatment at 1000 ºC. The SLM processed parts show a uniform fine equiaxed microstructure. The grain microstructure of SLM-fabricated specimens was much finer than that of the cast steel sample due to the high cooling rates associated with the AM process.

The grain microstructure of SLM-fabricated specimens was much finer than that of the cast steel sample due to the high cooling rates associated with the AM process. SEM micrograph revealed that annealing at a temperature 1000 ºC and after HIP treatment produced a uniform, fully recrystallized microstructure but with grains much finer than the cast material. It was reported that no difference in the peak patterns until annealing at 800 ºC. Further optimizations of the processing parameters are necessary for future work.AcknowledgmentThe authors acknowledge Dr. K. G. Prashanth at Leibniz Institute for Solid State and Materials Research Dresden (IFW) for producing SLM samples. Also, Mr. Brandon Creason at kittyhawkinc is highly thankful for performing the HIP treatments. One of the authors, Bandar AlMangour, would like to extend his appreciation to the Saudi Basic Industries Corporation, which generously awarded financial support to him.References[1] Wong KV, Hernandez A.

viewed PaperEffect of Powder Characteristics on Fatigue Performance of Additively Manufactured 17-4 PH Stainless SteelArun Poudel, Arash Soltani-Tehrani, Shuai Shao, Nima Shamsaei*National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849, USA Department of Mechanical Engineering, Auburn University, Auburn, AL 36849, USACorresponding author: shamsaei@auburn.eduAbstractThe characteristics of powder feedstock used during laser powder bed fusion (L-PBF) influence the mechanical performance of the fabricated parts. The flowability, spreadability, and internal porosity of the powder can affect the porosity formation and thus, impact the fatigue performances. Two batches of 17-4 precipitation hardening stainless steel powders supplied by two different vendors were used to fabricate the L-PBF specimens and investigate the effect of powder characteristics on porosity and fatigue behavior.

The powder batch with a wider particle size distribution, higher compressibility, higher cohesion between powder particles, and internal porosity resulted in a higher defect content in the fabricated specimens. Higher defect content led to inferior fatigue resistance along with more scatter in the fatigue lives. Fractography revealed the fatigue crack initiation from spherical pores as well as the lack of fusions in both batches.Keywords: Additive manufacturing, Powder characteristics, Particle size distribution, Volumetric defects, Fatigue performanceIntroductionDifferent additive manufacturing (AM) technologies such as laser powder bed fusion (L- PBF), laser powder directed energy deposition (LP-DED), electron beam directed energy deposition (EB-DED), etc., use metallic powder as feedstock during the fabrication process [1–5].

The physical characteristics of the feedstock powder such as particle size distribution, sphericity, internal porosity, etc., have been observed to influence the powder behavior during the AM processes [6]. Any influence in the powder behavior such as its flowability, spreadability, packing state, etc., can influence the porosity formation in the fabricated parts and thus, impact the mechanical properties [7–13]. Few studies have attempted to investigate the effect of powder characteristics on mechanical performances, especially fatigue performances [7–14].One of the major characteristics influencing the powder’s rheological behavior, and therefore the build quality, is the particle size distribution (PSD). It signifies the relative amount of different-sized powder particles in any specific batch. An AlSi7Mg powder with narrower PSD was observed to have lower cohesion between powder particles leading to better flowability and thus, higher density L-PBF parts [15].

An AlSi7Mg powder with narrower PSD was observed to have lower cohesion between powder particles leading to better flowability and thus, higher density L-PBF parts [15]. On the other hand, iron powder with a wider PSD was observed to fabricate lower-density L-PBF parts [16]. The smaller iron powder particles were observed to agglomerate, resulting in a poor packing state during the L-PBF fabrication process. In addition, the powder batches with narrower PSD have been observed to possess a better packing427state resulting in lower compressibility [11,17]. In a powder recycling study of Ti-6Al-4V material, the recycled powder batch with narrower PSD was observed to have better flowability, resulting in fewer and smaller pores, and thus, better fatigue performance [11].Rheological properties of the powder are also influenced by the particle’s surface features such as satellites and the sphericity of the powder particles.

The presence of satellites has been observed to increase the friction between the powder particles, impeding the flow of Ti-6Al-4V and nickel-based alloy powder [18]. It resulted in the parts with a lower density as well as lower ductility. In addition, the same study reported higher flowability of the powder with higher sphericity [18].Although a few studies have been done to understand the effect of powder characteristics on mechanical performance, only limited research is available correlating its effect on the fatigue behavior of AM materials. Furthermore, America Makes and the American National Standards Institute (ANSI) have considered “flowability”, “spreadability”, “particle size and particle size distribution”, “particle morphology”, “hollow particles and hollow particles with entrapped gas”, and “metal powder specifications” as some of the major technical gaps (PM1, 4, 6, and 7) in the additive manufacturing standardization collaborative (AMSC) roadmap [19].

Hence, this study compared two different batches of 17-4 precipitation hardening (PH) stainless steel (SS) metallic powders obtained from different vendors and correlated the effect of powder characteristics on porosity formation as well as fatigue performance of L-PBF parts. The 17-4 PH SS alloy was selected due to its variety of applications in aerospace, oil refinery, and biomedical industries [20,21].Material and Methods17-4 PH SS argon-atomized powders supplied by Carpenter Technology (i.e., Batch 1) and EOS (i.e., Batch 2) were utilized to fabricate cylindrical bars in EOS M290, an L-PBF AM machine. Bars were fabricated in a vertical orientation using recommended process parameters. The recommended process parameters for 17-4 PH SS include 220 W laser power, 756 mm/s laser speed, 100 µm hatch distance, 40 µm layer thickness, and 67° inter-layer rotation angle. The chemical compositions of both powder batches as reported by the powder’s manufacturer are listed in Table 1.

The chemical compositions of both powder batches as reported by the powder’s manufacturer are listed in Table 1. After fabrication, the bars were heat treated to CA-H1025 schedule as seen in Fig. 1(a) (i.e., solution annealing at 1050 °C for ½ hour followed by air cooling and aging at 552 °C for 4 hours followed by air cooling) [22]. The heat-treated bars were machined according to ASTM E466 [23] to final geometry shown in Fig. 1(b) and selected randomly from the build plate for fatigue testing. Fully-reversed force-controlled fatigue tests were performed at 500 MPa, 600 MPa, and 800 MPa stress levels using an MTS servo-hydraulic load frame.After fatigue failure, fractography was performed to investigate the crack initiating defects using Zeiss Crossbeam 550 scanning electron microscope (SEM).

Rheological properties of powders were investigated using Freeman Technology (FT4) powder rheometer according to ASTM D7891 [24], PSD using Anton Paar PSA 1190 based on laser diffraction technology [25], particle surface morphology using SEM, and internal porosity using Zeiss Xradia 620 Versa X- ray computed tomography (XCT) machine. The XCT scans were post-processed using ImageJ and Dragonfly Pro software [26,27]. In addition, the porosity in the fabricated specimens was investigated using Keyence VHX-6000 digital optical microscope. The porosity was investigated428in 4 different cross-sections (i.e., perpendicular to the building direction) throughout the gage- section of the specimens.Table 1 Chemical compositions of the 17-4 PH SS powders reported by the powder manufacturers.Element (Wt.

%) C Cr Ni Cu Mn Si Nb Mo N O P S FeBatch 1 0.010 16.300 4.180 4.100 0.220 0.400 0.250 0.020 0.040 0.050 0.012 0.004 Bal.Batch 2 0.010 16.670 4.270 3.680 0.050 0.020 0.310 0.080 0.100 0.030 0.010 0.004 Bal.Fig. 1 (a) CA-H1025 heat treatment schedule applied to 17-4 PH SS specimens. (b) The geometry of machined force-controlled fatigue specimens in accordance with ASTM E466 [23]. All dimensions are in mm.Results and DiscussionPowder CharacteristicsVarious characteristics of the feedstock powder such as its shape, size distribution, internal porosities, etc. influencing the flowability, spreadability, packing state, agglomerate formation, etc. were investigated. The surface morphology of the powder particles representative of the considered powder batches was examined and shown in Fig. 2. In general, the average particle size of Batch 2 was higher as compared to Batch 1.

2. In general, the average particle size of Batch 2 was higher as compared to Batch 1. Anomalies such as satellites were observed in both batches; however, it seemed to be higher for Batch 1(see Fig. 2(b & d)). The presence of these satellites can increase interparticle friction and result in lower flowability and spreadability [28,29]. Fig. 2(b) shows the presence of open pores in Batch 1 powder. Regarding sphericity, it429seemed to be higher in Batch 2 as compared to Batch 1 (see Figs. 2(b & d)); however, a thorough analysis needed to be performed to support the statement. Higher sphericity has been observed to increase the spreadability of the powder particles [18].Fig. 2 Particle morphology of (a & b) Batch 1 and (c & d) Batch 2 powder. Red and blue arrows point to satellites and open pores in the powder particles, respectively.The PSD of each powder batch was investigated three times and their average is shown in Fig. 3.

Red and blue arrows point to satellites and open pores in the powder particles, respectively.The PSD of each powder batch was investigated three times and their average is shown in Fig. 3. The average particle size in Batch 2 was observed to be larger than in Batch 1. In addition, Batch 1 contained a larger number of smaller particles as compared to Batch 2. The smaller powder particles tend to adhere with each other and also with larger particles, and form agglomerates [9,16]. Such agglomerates can result in lower flowability as well as a poor packing state of the powder due to the presence of air/void between them. Fig. 3(c) presents the D10, D50, and D90 measures of both powder batches, where Dx signifies that x% of the powder particles is below its value. The D10, D50, and D90 for Batch 2 were 40%, 43%, and 38% higher as compared to Batch 1. As seen in the plot, the D10s of Batch 1 and Batch 2 are smaller than 21.7 µm and 30.4 µm, respectively.

The D10, D50, and D90 for Batch 2 were 40%, 43%, and 38% higher as compared to Batch 1. As seen in the plot, the D10s of Batch 1 and Batch 2 are smaller than 21.7 µm and 30.4 µm, respectively. This further validates the statement that Batch 1 is more prone to the formation of agglomerates as compared to Batch 2. From the previous study, the compressibility—which signifies the amount of air in between the powder particles—of Batch 1 was 67% higher than Batch 2 [14]. In addition, a good correlation was observed between the shear stress required for the powder batch to flow at different applied normal stresses. Batch 1 required higher shear stress to flow as compared to Batch 2 suggesting higher friction or resistance in between the Batch 1 powder particles.

Batch 1 required higher shear stress to flow as compared to Batch 2 suggesting higher friction or resistance in between the Batch 1 powder particles. This behavior was attributed to the higher cohesion (i.e., two times) in between Batch 1 powder particles as compared to Batch 2 [14].The span (i.e., span = (D90-D10)/D50) was calculated to be 0.8 and 0.76 for Batch 1 and Batch 2, respectively [29]. The lower span of Batch 2 indicates a narrower PSD and a more uniform particle size. Typically, fine particles attaching to larger particles have been observed to degrade powder's flowability and spreadability characteristics. Hence, it can be expected that the Batch 2 powder particles flow more smoothly and uniformly across the build plate as compared to Batch 1.430Fig. 3 Comparison of PSD between powder batches: (a) Volumetric probabilistic density distribution, (b) cumulative density distribution, and (c) D10, D50, and D90 measurements.

3 Comparison of PSD between powder batches: (a) Volumetric probabilistic density distribution, (b) cumulative density distribution, and (c) D10, D50, and D90 measurements. (d) The comparison of the shear stress required for the powders to flow at different stress levels.Porosity in Powder Particles and Fabricated SpecimensThe XCT findings of both powder batches are illustrated in Fig. 4. The voxel size of 0.73 µm was used during the XCT scans of 0.3055 mm3 volume. In general, Batch 1 contains smaller but more internal pores as compared to Batch 2. A few large-size internal pores were evident in Batch 2 as seen in Figs. 4(d-f). The largest pores inside the powder particles in both batches were in the range of 30 µm to 33 µm.

A few large-size internal pores were evident in Batch 2 as seen in Figs. 4(d-f). The largest pores inside the powder particles in both batches were in the range of 30 µm to 33 µm. During fabrication, these entrapped pores inside the powder particles may escape from the melt-pool; however, if the Marangoni force—which can push the gas bubbles downward—is higher than the buoyancy force, these particles get trapped in the substrate during the solidification process [30]. The presence of a large number of internal pores in Batch 1 powder particles as well as its poorer packing state made it susceptible to the formation of a higher degree of volumetric defects in the fabricated parts [31–33]. The statistics of the intra- particle pores is presented in Fig. 5(a). As shown, the pore count in Batch 1 powder was significantly higher than in Batch 2 for most size ranges. During the analysis, defects with sizes smaller than 10 µm were discarded.431Fig.

5(a). As shown, the pore count in Batch 1 powder was significantly higher than in Batch 2 for most size ranges. During the analysis, defects with sizes smaller than 10 µm were discarded.431Fig. 4 Internal porosities observed in (a & b) Batch 1 and (d & e) Batch 2 powder particles. The color scale signifies the equivalent diameter of internal pores in the µm scale. In addition, one volumetric scan slice from (c) Batch 1 and (f) Batch 2 is presented.After fabrication, porosity analysis of specimens was performed using digital optical microscopy to investigate the influence of powder characteristics on defect content (see Fig. 5(b)). Parts fabricated using Batch 1 resulted in more volumetric defects for all size ranges. The size range of the defects observed in both powder particles and specimens was similar. In addition, the standard deviation observed during porosity analysis between 4 different layers was found to be higher for Batch 1 specimens. The observation made in Fig.

In addition, the standard deviation observed during porosity analysis between 4 different layers was found to be higher for Batch 1 specimens. The observation made in Fig. 5(b) agrees with the earlier findings regarding flowability, packing state, as well as internal porosity in the powder particles. The presence of agglomerates and poor spreadability across the build plate for Batch 1 might have resulted in a non-uniform powder layer during the fabrication process. In addition, the presence of a higher degree of intra-particle porosity in Batch 1 (see Fig. 4) exacerbated the process and resulted in a higher porosity.432Fig. 5 Size distributions of defects observed in (a) particles of the two powder batches via XCT and (b) fabricated specimens via digital optical microscopy.Fatigue PerformanceThe stress-life fatigue behavior of specimens fabricated using both powder batches is shown in Fig. 6. The fatigue behavior of Batch 1 specimens was adapted from Soltani-Tehrani et al.

6. The fatigue behavior of Batch 1 specimens was adapted from Soltani-Tehrani et al. [9]. The fatigue resistance of both powder batches was comparable at low cycle fatigue (LCF) regime. It was attributed to the similar microstructure in both materials after being subjected to the same, CA-H1025 heat treatment schedule. However, the difference was more prominent in mid- cycle fatigue (MCF) and high cycle fatigue (HCF) regimes. The fatigue lives observed for Batch 2 specimens were higher in MCF and HCF regimes as compared to Batch 1. Furthermore, one Batch 2 specimen tested at 500 MPa was observed to be run out (i.e., exceeded five million cycles). Since the fatigue resistance of the material is more sensitive to the volumetric defects in HCF, the scatter was also observed to be higher at HCF for both powder batches. The higher level of porosity observed in Batch 1 specimens as compared to Batch 2 (see Fig. 5(b)) agrees with the fatigue performances observed in Fig.

The higher level of porosity observed in Batch 1 specimens as compared to Batch 2 (see Fig. 5(b)) agrees with the fatigue performances observed in Fig. 6.Fractography was performed to examine the crack initiating defects in both batches of specimens. Since the fatigue specimens were machined and polished to a mirror finish, the effect of surface roughness was eradicated and the cracks were observed to have initiated from the volumetric defects. In both batches, cracks initiated from spherical defects as well as irregular- shaped lack of fusions. Closer inspection revealed that the cracks in a few of the Batch 1 specimens initiated from more than one defect grouped together as seen in Fig. 7(a). Effectively, the closely grouped, smaller defects behaved as one big defect [9].433Fig. 6 Stress-life fatigue behavior of L-PBF CA-H1025 17-4 PH SS machined specimens fabricated using Batch 1 and Batch 2 powder batches.Fig.

6 Stress-life fatigue behavior of L-PBF CA-H1025 17-4 PH SS machined specimens fabricated using Batch 1 and Batch 2 powder batches.Fig. 7 Fractography of L-PBF CA-H1025 17-4 PH SS machined specimens fabricated using (a) Batch 1 and (b) Batch 2.ConclusionsCharacteristics of powder feedstocks such as particle morphology, particle size distribution, internal porosity, etc., were observed to influence their flowability, spreadability, packing state, etc. This resulted in different degrees of porosity in the fabricated specimens and thus, different fatigue performances. The conclusions drawn from this study are:The presence of finer powder particles and higher cohesive force between them resulted in the formation of agglomerates, and thus led to a poor packing state.Lower flowability and poor packing state of Batch 1 powder resulted in higher porosity in the fabricated specimens.Higher porosity deteriorated the fatigue resistance of Batch 1 specimens.

In addition, it also resulted in more scatter in the fatigue lives of Batch 1 specimens as compared to Batch 2. • Fatigue cracks were observed to initiate from spherical and irregular-shaped lack of fusion defects.434AcknowledgmentsThis material is based upon the work partially supported by National Science Foundation (NSF) under grant # 1919818.References[1] W.E. Frazier, Metal additive manufacturing: A review, J. Mater. Eng. Perform. 23 (2014)1917–1928. https://doi.org/10.1007/s11665-014-0958-z.[2] A. Cooke, J. Slotwinski, Properties of Metal Powders for Additive Manufacturing: A the Art of Metal Powder Property Testing, (n.d.).Review of https://doi.org/10.6028/NIST.IR.7873.the State of[3] R. Ghiaasiaan, A. Poudel, N. Ahmad, P.R. Gradl, S. Shao, N. Shamsaei, High Temperature Tensile and Fatigue Behaviors of Additively Manufactured IN625 and IN718, Procedia Struct. Integr. 38 (2022) 581–587. https://doi.org/10.1016/J.PROSTR.2022.03.059. [4] R. Ghiaasiaan, A.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of powder characteristics on tensile properties of additivelymanufactured 17-4 PH stainless steelArun Poudel, Arash Soltani-Tehrani, Shuai Shao, Nima Shamsaei*National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849, USADepartment of Mechanical Engineering, Auburn University, Auburn, AL 36849, USACorresponding author: shamsaei@auburn.eduAbstractLaser beam powder bed fusion (LB-PBF) process uses metallic powders as feedstock, whose particle characteristics such as cohesion, compressibility, size distribution, etc., can vary and affect the mechanical performance of the fabricated parts.

In this study, two powder batches of 17-4 precipitation hardening (PH) stainless steel (SS) supplied by EOS (Batch 1) and Carpenter Technology (Batch 2) were used to fabricate specimens using identical process parameters to understand the effects of particle characteristics on defect content as well as tensile performance of the LB-PBF specimens. Higher cohesion and compressibility as well as lower sphericity in Batch 2 resulted in specimens with higher porosity levels. During tensile testing, the higher porosity level in Batch 2 yielded lower ductility.

During tensile testing, the higher porosity level in Batch 2 yielded lower ductility. In contrast, the microstructure was observed to be less sensitive to particle characteristics because of which the tensile strengths of the specimens were found to be comparable to each other.KEYWORDS: Additive manufacturing, Tensile performance, Powder characteristics, Powder flowability, Porosity, LB-PBFIntroductionOne of the common techniques to generate powder is gas atomization, and the output can be used in powder-based additive manufacturing (AM) technologies such as laser or electron beam powder bed fusion (LB- or EB-PBF) [1]. Therefore, the rheological behavior of the gas-atomized powders such as their flowability, spreadability, and packing density can influence the mechanical performance of the resulting parts.

Therefore, the rheological behavior of the gas-atomized powders such as their flowability, spreadability, and packing density can influence the mechanical performance of the resulting parts. Powder rheological behavior is itself affected by many factors including but not limited to particle size distribution (PSD), cohesion, internal porosity, and sphericity of the powder particles. Studies attempting to understand the influence of these powder characteristics on the mechanical performances of the additively manufactured (AMed) parts are relatively limited [2–4].Powder characteristics such as sphericity and PSD are a few of the most influential characteristics of powder flowability. Higher sphericity, as well as fewer satellites (i.e., smaller powder particles being attached to the bigger ones [5]), can typically result in higher flowability and consequently denser LB-PBF parts [4,6,7]. Muniz-Lerma et al.

Muniz-Lerma et al. [8] also reported that a wider PSD can result in more agglomerates and higher cohesion between the powder particles and870ultimately less dense materials. Similar observations were reported by Simchi [9] that the existence of very fine and coarse/large particles within the powder batch (i.e. a wider PSD) can adversely affect the part density. The reason was explained by the tendency of fine particles to agglomerate and large particles to segregate.Comparable observations were also reported by Carrion et al. [10] and Soltani-Tehrani et al. [11]. In these studies, the narrower PSD which included the particles with more uniform sizes was seen to result in not only a higher powder flowability but also a higher packing density. It was detailed that the existence of agglomerates within the batch can lead to more empty spaces within the powder bulk.

It was detailed that the existence of agglomerates within the batch can lead to more empty spaces within the powder bulk. These empty spaces were reported to negatively influence the consolidation of the powder layers to the previously solidified layers during the LB-PBF process. As a result, more pores with larger sizes were observed in the LB-PBF parts fabricated from powder batches with wider PSDs.Although it is well established that particle morphology and PSD can affect powder behavior and consequently mechanical properties, limited research studies have correlated such variations in powder with the AMed part performance. As a result, this study aims to provide some insights into the relationships between the powder feedstock and structure-part performance.

As a result, this study aims to provide some insights into the relationships between the powder feedstock and structure-part performance. It is also noteworthy to mention that America Makes and the American National Standards Institute (ANSI) have reported the “flowability”, “spreadability”, “particle size and particle size distribution”, “particle morphology”, “hollow particles and hollow particles with entrapped gas”, and “metal powder specifications” as some major technical gaps (PM1, 4, 6, and 7) in the additive manufacturing standardization collaborative (AMSC) roadmap [12], necessitating more in-depth investigations and further research.Material and methodsTwo different batches of argon atomized 17-4 PH stainless steel (SS) powder supplied by EOS (Batch 1) and Carpenter Technology (Batch 2) were used to fabricate some cylindrical bars with an 11 mm diameter and 84 mm length using identical process parameters in the EOS M290 LB-PBF AM machine.

The chemical composition of each of the powder batches, reported by the manufacturer, is shown in Table 1. EOS has reported the chemical composition of its powder following the ASTM F899 standard. After fabrication, the cylindrical bars were subjected to the CA-H1025 heat treatment procedure (i.e. solution annealing at 1050 °C for ½ hour followed by air cooling and aging at 552 °C for 4 hours and the subsequent air cooling) to homogenize and strengthen the microstructure [13]. Finally, the heat-treated specimens were machined into the net- shape fatigue specimens with tangentially blending fillets between the test section and the ends following ASTM E466 for uniaxial force-controlled fatigue testing [14]. The tensile tests were performed using the fatigue specimens on an MTS servo-hydraulic load frame with a 25 KN load cell according to ASTM E8 standards [15].871Fig. 1.

The tensile tests were performed using the fatigue specimens on an MTS servo-hydraulic load frame with a 25 KN load cell according to ASTM E8 standards [15].871Fig. 1. The geometry of the machined specimens following ASTM E466 standards [16].Table 1 Chemical composition of the 17-4 PH SS powder reported by the powder’smanufacturers (i.e., Batch 1 from EOS and Batch 2 from Carpenter technology).Element (Wt. %) C Cr Ni Cu Mn Si Nb Mo N O P S FeBatch 1 0.010 16.300 4.180 4.100 0.220 0.400 0.250 0.020 0.040 0.050 0.012 0.004 Bal.Batch 2 0.010 16.670 4.270 3.680 0.050 0.020 0.310 0.080 0.100 0.030 0.010 0.004 Bal.Powder characteristics such as compressibility, cohesion, bulk and tapped densities, and shear stress (according to ASTM D7891 [17]) of both powder batches were investigated using Freeman Technology (FT4) powder rheometer. Furthermore, the PSD was analyzed as per ASTM E2651 standard [18] using an Anton Paar PSA 1190 with laser diffraction technology.

Furthermore, the PSD was analyzed as per ASTM E2651 standard [18] using an Anton Paar PSA 1190 with laser diffraction technology. The internal porosity of the powders particles and their morphology was studied using a Zeiss Xradia 620 Versa, an X-ray computed tomography (CT) machine and a Zeiss Crossbeam 550 SEM/FIB scanning electron microscope (SEM), respectively. In addition, the porosity content in the fabricated specimens was studied and quantified on polished cross sections by using Keyence VHX-6000 digital optical microscope.Results and discussionPowder characteristicsDifferent characteristics of the powder such as its flowability, spreadability, packing density can influence the part performances. Even though the parameters such as flowability and872spreadability represent the powder flow throughout the build plate, they can also have a major influence on the porosity formation of the resulting AM parts. As seen in Figs.

As seen in Figs. 2 (a & b), the powder particles’ morphology illustrated the existence of both spherical particles and agglomerates in both powder batches. However, visually comparing both batches, Batch 2 seems to consist of particles with relatively high sphericity as can be seen in Fig. 2(b) while the particles in Batch 1 have more irregularly-shaped particles (see Fig. 2(a)). In addition, comparing the SEM images, Batch 1 had a higher level of open porosity in the powder particles as compared to Batch 2. The analysis performed using SEM agrees with the findings from the X-ray CT scanning of the powders in Figs. 2 (c & d). As seen in these figures, the level of internal porosity (colored in red) in the Batch 1 powder particles is higher than in Batch 2. Some studies have reported that the existence of internal particle porosity contributes to the gas entrapment and less dense LB-PBF fabricated parts [2,19,20].(c)(d)Fig. 2.

Some studies have reported that the existence of internal particle porosity contributes to the gas entrapment and less dense LB-PBF fabricated parts [2,19,20].(c)(d)Fig. 2. Morphology of (a) Batch 1, and (b) Batch 2 17-4 PH SS powder particles obtained via Zeiss Crossbeam 550 SEM, and X-ray CT 3D visualizations of (c) Batch 1 and (b) Batch 2 powders illustrating internal porosities and morphology.Also, in Fig. 3(a), the PSD of both powder batches illustrates a clear difference in the mean particle diameter and also the existence of finer particles in Batch 1. The mean particle size of Batch 2 was observed to be 49 µm which is 40% larger than that of Batch 1. Similarly, the D10 and873D90 of the Batch 1 powder was found to be 22 µm and 47 µm whereas for Batch 2, it was 30 µm and 65 µm respectively. Therefore, it can be distinctly concluded that more fine particles exist in Batch 1.

Therefore, it can be distinctly concluded that more fine particles exist in Batch 1. Fine powder particles have been observed to decrease the powder flowability and resulting in inferior powder spreadability on the build plate during fabrication due to higher interparticle friction [21,22].This observation also supports the finding regarding the densities of powder particles in both apparent and tapped conditions. The Hausner ratio (i.e., the ratio of tapped to apparent powder density) of Batch 1 and 2 was found to be 1.08 and 1.04, respectively. Both apparent and tapped densities were observed to be lower in Batch 1 because of the higher internal porosity and tendency to form agglomerates. The non-uniform packing state caused by the formation of the agglomerates can result in a higher percentage of compressibility due to more void spaces among particles, and also higher cohesion (see Fig. 3 (c & d)).

3 (c & d)). The higher cohesion in Batch1 due to the presence of more fine particles can hinder the powder flow and its uniform spreadability across the build plate [8].4.75Batch 14.5Batch 2) lm / g ( y t i s n e D4.2543.753.5Apparent 1(b)Tapped 2(a)(c)(d)Fig. 3. Comparison of the 17-4 PH SS powders’ characteristics: (a) PSD, (b) apparent and tapped densities, (c) compressibility, and (d) cohesion.874Porosity level of fabricated partsPorosity analysis of the specimens fabricated using Batches 1 and 2 powders with their corresponding pore size distribution and its standard deviations between 4 layers is shown in Fig. 4. It seems that the specimens fabricated using Batch 2 powder have a lower porosity level. This behavior was also consistent with the rheological findings where Batch 2 was observed to have a higher flowability and superior packing state due to lower cohesivity, narrower PSD, less internal particle porosity, and higher sphericity.

In addition, the standard deviation in the porosity count in the specimen fabricated using Batch 2 powder is much lower as compared to Batch 1 in all porosity size ranges which can suggest the consistency in the porosity distribution between consecutive layers. Similar findings have been reported in Ref. [2] where the powder with finer particles resulted in more agglomerates, a non-uniform powder distribution, and less dense materials.Fig. 4. Comparison between the porosity levels of LB-PBF 17-4 PH SS specimens fabricated from Batches 1 and 2.Tensile propertiesThe quasi-static tensile behaviors of LB-PBF 17-4 PH SS parts fabricated from Batch 1 and 2 are presented in Fig. 5. The ultimate tensile strength (UTS) and yield strength (YS) of Batch 1 specimens were slightly higher than Batch 2 (~7%). This small difference in tensile strength can be attributed to the different powder densities. As noted in Fig. 3(b), the bulk density of Batch 2 was 8.5% higher than Batch 1.

This small difference in tensile strength can be attributed to the different powder densities. As noted in Fig. 3(b), the bulk density of Batch 2 was 8.5% higher than Batch 1. Typically, powder batches with smaller particle sizes should have better thermal contact and better thermal conductivity. Therefore, it can be assumed that the higher thermal conductivity of Batch 1 has resulted in more heat dissipation toward the build plate, higher temperature gradient in the part, and consequently higher cooling rates during fabrication. As a result, the grains are expected to be somewhat larger in Batch 2 specimens [23,24]. Similar values of UTS and YS were also reported by Nezhadfar et al. [25] for 17-4 PH SS material subjected to CA-H1025 heat-treatment process. In terms of ductility, the strain to fracture (ɛf) for Batch 2 was observed to be 8.6% higher than Batch 1.

[25] for 17-4 PH SS material subjected to CA-H1025 heat-treatment process. In terms of ductility, the strain to fracture (ɛf) for Batch 2 was observed to be 8.6% higher than Batch 1. Since the defect content in the specimens can influence the ductility of the materials, the lower porosity level in Batch 2 should have contributed to its875superior ductility. Similar observations were reported in Ref. [4], where a lower ductility was correlated with higher porosity resulted from lower flowability as well as more fine particles within the batch. The comparison of the tensile properties of the 17-4 PH SS material obtained in this study with the ones in the literature can be seen in Table 2.Fig. 5. Comparison between the UTS, YS and ɛf of the LB-PBF 17-4 PH SS specimens fabricated using Batch 1 and 2 powders.Table 2 Comparison of tensile strengths obtained in this study with the ones from theliterature.17-4 PH SSBatch 1 Batch 2 Nezhadfar et al.

[25] Carpenter Technology [26] Wrought (EOS) [27]Heat TreatmentCA-H1025 CA-H1025 CA-H1025 H950 CA-H900UTS (MPa) 1203 1119 1167 1172 1336YS (MPa) 1155 1085 1140 1069 1235Strain to fracture (mm/mm) 0.35 0.38 0.30 N/A N/AConclusionsThe powder characteristics such as compressibility, cohesion, morphology, PSD, etc. were observed to influence the porosity as well as the mechanical properties of the AMed 17-4 PH SS parts.Higher compressibility, cohesion, and the presence of more irregularly shaped and fine particles in Batch 1 resulted in higher porosities.Ductility was lower in the Batch 1 specimens due to the higher level of porosity.876As opposed to ductility, the tensile strength was observed to be less sensitive to the variations in powder properties.AcknowledgmentsThis material is also based upon work partially supported by the National ScienceFoundation (NSF) under grant # 1919818.References[1]P. Moghimian, T. Poirié, M. Habibnejad-Korayem, J.A.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEFFECT OF POWDER CHEMICAL COMPOSITION ON MICROSTRUCTURES AND MECHANICAL PROPERTIES OF L-PBF PROCESSED 17-4 PH STAINLESS STEEL IN THE AS-BUILT AND HARDENED-H900 CONDITIONSS. Vunnam*, S. Dobson*, A. Saboo†, D. Frankel†, C. Sudbrack†, and T. L. Starr *J. B. Speed School of Engineering, University of Louisville, Louisville, KY, 40292, USA †QuesTek Innovations LLC, 1820 Ridge Avenue, Evanston, IL USA 60201, USAAbstractPost-build heat treatments such as solutionizing and precipitation hardening are recommended for selective to achieve a homogeneous microstructure. In this study, the effect of powder elemental composition on microstructures and mechanical properties of SLM processed 17-4 PH was studied in the as-built and precipitation hardened (H900) condition without prior solutionizing.

Microstructural characterization demonstrated that H900 increased martensite phase composition for samples from powder with low chromium to nickel equivalent (Creq/Nieq) value, whereas no significant difference was observed for the samples from powders with high Creq/Nieq value. None of the specimens exhibited austenite reversion and strain hardening behavior in the as-built and H900 conditions. Low Creq/Nieq specimen exhibited higher yield and tensile strengths, and microhardness from H900, which are comparable to H900 wrought sample.

Low Creq/Nieq specimen exhibited higher yield and tensile strengths, and microhardness from H900, which are comparable to H900 wrought sample. However, no significant improvement in total elongation was observed other than uniform elongation for low Creq/Nieq specimen.laser melting (SLM) processed componentsKeywords: 17-4 PH stainless steel, Selective laser melting, Electron backscatter diffraction (EBSD), Phase composition, Mechanical Properties, H900 hardeningIntroductionLaser powder bed fusion (L-PBF) processing of precipitation hardenable (PH) martensitic stainless steels, specifically 17-4 PH, are of great interest because of their unique combination of high strength and corrosion resistance. The microstructure and mechanical properties of L-PBF processed 17-4 PH stainless steel alloy parts are highly dependent on the powder chemical composition and the post-build heat treatments.

The microstructure and mechanical properties of L-PBF processed 17-4 PH stainless steel alloy parts are highly dependent on the powder chemical composition and the post-build heat treatments. Recent studies focused on improving yield strength, hardness and fatigue behavior by optimizing processing conditions and heat treatments [1-7]. State of the art literature on L-PBF processed 17-4 PH reveals a strong dependence of material microstructure and phase constitution on process parameters, particularly the presence of a signiﬁcant fraction of retained austenite [1, 3, 4, 8] or retained δ ferrite [9] in the as-built condition. Extremely high temperature gradients induce nonhomogeneous grain size and residual stresses in L-PBF parts [10, 11]. Post-build heat treatments were recommended to relieve the stress and to obtain uniform grain size.

Post-build heat treatments were recommended to relieve the stress and to obtain uniform grain size. Solutionizing is used for uniform distribution of the elements in the alloy, whereas precipitation hardening induces nanoscale Cu rich precipitates which play a key role in enhancing the mechanical strength and hardness. However, heat treatments are time consuming and expensive. Very limited research has been focused on microstructure and mechanical properties in the as-built condition as well as H900 condition without prior solutionizing. The effect of atomizing gas on microstructure and mechanical properties in the as- built and hardened condition was studied [12]. Hardening using H900 alone without solutionizing inhibited strengthening due to the large volume of retained austenite [13]. However, the alloy463powder chemical composition effect on microstructure and tensile properties was not explored in detail.

However, the alloy463powder chemical composition effect on microstructure and tensile properties was not explored in detail. The objective of the current study is to determine the effect of powder chemical composition on the microstructure and mechanical properties of L-PBF processed 17-4 PH in the as-built and H900 conditions. Previous work by the authors of this paper outlined the L-PBF processed 17-4 PH texture and phase composition sensitivity to powder chemical composition in the as-built condition [14]. Powder with lower Creq/Nieq value produced a component with largely martensitic phase with some residual δ ferrite and retained austenite. In this study, microstructure and mechanical properties of components from three different commercially available Ar-atomized 17-4 PH powders that conforms to the UNS S17400 specification range were explored in the as- built and H900 conditions and compared with wrought under identical conditions.

In-depth microstructure characterization was carried out using optical microscopy (OM) and electron backscatter diffraction (EBSD) techniques. Microhardness and tensile properties were measured to validate the microstructural analysis.ExperimentalIn this study, 17-4 PH stainless steel test specimens were built using a commercial EOSINT M290 SLM machine. Three commercially available Ar-atomized powders (Powder A, Powder B and Powder C) with a different chemical composition that conforms to the UNS S17400 specification range were used for this study. The chemical composition of wrought 17-4 PH and the three powders and Creq/Nieq values based on WRC-1992 equations were tabulated in Table 1 [15]. All the test specimens were fabricated with the optimized SLM parameters using a standard build strategy, which involves bidirectional beam scanning with a scan rotation angle of 67o, and all the parts exhibited a full density.

Detailed powder properties and part densities were discussed elsewhere [16]. Processing parameters deployed in this study included laser power of 190 W, scanning speed of 750 mm/s, hatch distance of 0.100 mm, layer thickness of 0.040 mm and build plate temperature of 40 °C. As-received wrought material and the as-built SLM test specimens were subjected to H900 heat treatment at 482 oC for 1 hour then air cooled. For shorthand notation, “AB” and “H900” suffixes attached to the name of the powder for the as-built and H900 conditions, respectively.As-received (annealed) wrought samples (Wrought-AR) and H900 wrought samples (Wrought- H900) were cut such that the section surface is normal to the rolling direction. Metallographic sample preparation was done using a specific metallographic sample preparation method that can be found elsewhere [14]. The freshly polished sample’s surface was etched by a Kalling’s no. 2 reagent for 1 to 2 minutes using an immersion method.

The freshly polished sample’s surface was etched by a Kalling’s no. 2 reagent for 1 to 2 minutes using an immersion method. The etched samples were rinsed in isopropyl alcohol immediately and air-dried. The etched samples were examined along the build direction using an Olympus BX53 optical microscope. Samples were further polished using GIGA 0900 vibratory polisher using colloidal silica suspension for EBSD. EBSD was carried out at 20 kV using an FEI Teneo field emission SEM equipped with an EDAX TEAM EBSD system. For EBSD analysis, the polished samples from sections parallel to build direction were tilted 70o to the electron beam and 12 mm working distance was maintained. The step size used for the EBSD analysis ranges from 0.3 to 1.5 μm. The EBSD data acquired were processed using TSL OIM analysis software.464Tensile testing was performed in an Instron 50 kN test machine (Model: 5569A) using ASTM E8-16 standard tensile testing method.

Tensile testing specimens were machined out using electrical discharge machining (EDM) from the bulk material primitives according to the dimensions specified by the ASTM E8 standard to mitigate the surface related artifacts of the SLM process. The hardness of the samples was evaluated by performing Vickers micro-hardness measurements on the polished sections using Shimadzu’s HMV-G21. At least 10 measurements were taken diagonally from each sample using a load of 980.67 mN (100 gf) with a 10 s holding time.Table 1.

At least 10 measurements were taken diagonally from each sample using a load of 980.67 mN (100 gf) with a 10 s holding time.Table 1. Chemical composition (wt.%) of the three Ar-atomized 17-4 PH stainless steel powders and as-received wroughtCrNi Mo Nb+TaCNCu MnSiFe Creq/NieqPowder A16.124.150.20.020.007 3.210.79 Bal.2.81Powder B16.74.29 0.050.270.010.031 4.490.20.36 Bal.2.65Powder C15.64.70.220.010.036 3.610.50.38 Bal.2.36Wrought 15-17.5 3-5 0-0.5 0-0.45 0-0.070-5.0 0-1.0 0-1.0 Bal.Provider datasheet *Creq = Cr + Mo + (0.7 x Nb) *Nieq = Ni + (35 x C) + (20 x N) + (0.25 x Cu)Results and DiscussionMicrostructure and Phase CompositionOptical micrographs of etched samples in the as-built and H900 conditions along the build direction are shown in Fig. 1. Powder A and Powder B samples have columnar grains that are grown across several melt pool boundaries along the build direction.

1. Powder A and Powder B samples have columnar grains that are grown across several melt pool boundaries along the build direction. Small martensitic laths can be seen below the melt pool boundary for Powder A sample. There is no significant difference observed in the microstructures of Powder A, Powder B and Powder C samples after H900 heat treatment from optical micrographs. Powder A and Powder B samples exhibited a large δ-ferrite phase constituent after H900 heat treatment similar to the as-built samples. However, Powder C sample has a fine grain microstructure in the as-built as well as in heated conditions. In order to get a more detailed understanding of Powder C sample microstructure, further analysis was performed using EBSD and compared with the Wrought-H900 sample. In-depth microstructure and phase evolution analysis for all three SLM samples in the as-built condition can be found elsewhere [14].

In-depth microstructure and phase evolution analysis for all three SLM samples in the as-built condition can be found elsewhere [14]. EBSD orientation maps obtained from an area of 400 μm x 300 μm for Powder C- H900 and Wrought-H900 are shown in Fig.2(a,b). FCC-austenite grains can be clearly seen in465higher magnification orientation maps obtained from an area of 99 μm x 72 μm as shown in Fig. 2(c, d) (separated by a black line). EBSD phase analysis indicates that austenite volume fraction in Powder C-H900 and Wrought-H900 samples are 2.3 and 1.2%, respectively. Powder C-H900 sample exhibited inhomogeneous grain size distribution by having coarse columnar grains whereas, Wrought-H900 has more homogeneous grain size distribution. The average grain size for Powder C-H900 sample is 9.5 μm including FCC and BCC grains by excluding low angle grain boundaries (LAGBS) for martensitic laths.

The average grain size for Powder C-H900 sample is 9.5 μm including FCC and BCC grains by excluding low angle grain boundaries (LAGBS) for martensitic laths. The average martensitic block size for Powder C-H900 and Wrought-H900 samples are 12.6 and 4.3 μm, respectively. Powder C-H900 sample has fine grain regions from transformed martensite whereas coarse columnar grains are from retained δ ferrite [14]. Fig. 3 shows overlay of orientation and image quality (IQ) maps for same maps from Fig. 2(c, d). It can be seen the darker contrast within the grains in martensitic region due to the poor image quality from the distorted BCT crystal structure. EBSD grain boundary maps corresponding to orientation maps in Fig. 2(c, d) and 3 (a, b) are shown in Fig. 3(c, d). Grain boundaries are color-coded by green (2-17o), red (17-49o) and black (49-180o).

EBSD grain boundary maps corresponding to orientation maps in Fig. 2(c, d) and 3 (a, b) are shown in Fig. 3(c, d). Grain boundaries are color-coded by green (2-17o), red (17-49o) and black (49-180o). LAGBs (2-17o) were observed in the fine grain region but not in the columnar grains for Powder C-H900 sample indicating that columnar grains are δ ferrite grains and were not transformed to martensite upon cooling. The linear fraction of LAGBs for Powder-C sample is 48.1% and increased from 35% in the as-built condition [14]. The linear fraction of Wrought-H900 sample is 59.6% and these are uniformly distributed throughout the specimen. As shown in Fig. 4, frequency of martensitic boundaries observed above 49o and below 17o. Parent austenite grain boundaries and lath boundaries (specifically wrought sample) can be seen in Fig. 3(a, b) which arise from the IQ contrast. As shown in Fig.

2(c, d)) are shown in Fig. 5. Both Powder C-H900 and Wrought-H900 samples have higher {100} texture strength compared to {110} and {111}. Weaker {100} texture strength for Powder C-H900 sample compared to Wrought-H900 sample could be due to the large grain growth rate, which led to parent austenite grain refinement which can be seen in Fig. 3(c). Stronger texture for wrought sample can be due to large plastic deformation and subsequent recrystallization from hot rolling as the sample surface is normal to the rolling direction [17].468Fig. 3. EBSD IPF and IQ overlay maps (a, b), grain boundary misorientation maps: 2°– 17°(green), 17°–49°(red), and 49°–180°(black)46915Wrought-H900 Powder C-H900)%10(y c n e u q e r F501020 50 Misorientation angle(o)30406070Fig. 4. The frequency distribution of parent austenite grain boundaries vs.

4. The frequency distribution of parent austenite grain boundaries vs. misorientation anglePowder C-H900 A1max= 3.393 2.768 2.258 1.842 1.503 1.226 1.000 0.816Wrought-H900A1max= 4.600 3.567 2.766 2.145 1.663 1.290 1.000 0.775001001Fig. 5. (100) pole figures along build direction for Powder C-H900 and Wrought H900 samplesFig. 6(a, b) shows the XRD spectra that were obtained from the specimen sections along the build direction for all three Ar atomized powder and wrought samples. XRD scans from 40o to 50o470diffraction angle with increased step size (2s/step) were obtained to confirm that no low-intensity austenite peak at 43.5o is present. From XRD data no detectible metastable austenite is present in any of these samples. Table 2 details the volume fraction of austenite observed for all the samples from XRD and EBSD analysis.

From XRD data no detectible metastable austenite is present in any of these samples. Table 2 details the volume fraction of austenite observed for all the samples from XRD and EBSD analysis. EBSD phase analysis exhibited austenite volume fractions for Powder C-H900 and Wrought-H900 samples are approximately 2.3 and 1.2%, respectively. In general, the standard heat treatment process for 17-4 PH [13, 18, 19] causes austenite reversion from martensite due to the chemical stabilization by the diffusion of the alloying elements. However, XRD and EBSD results indicate that there is no significant austenite reversion occurred in Powder C and wrought samples from H900 heat treatment. All three SLM processed samples have no significant amount of austenite in the as-built condition as well [14].

All three SLM processed samples have no significant amount of austenite in the as-built condition as well [14]. Out of various possibilities, low concentration of austenite stabilizing elements specifically Ni and Mn in Powder A and Powder B, high cooling rate from SLM fabrication process accompanied by no additional austenite stabilizer such as N2 gas from powder fabrication process could suppress the austenite transformation from δ ferrite upon cooling. More interestingly, EBSD results exhibited lower austenite volume fraction in Powder C-H900 sample compared to the as-built sample indicating that there is no austenite reversion from hardening. XRD patterns also indicate that Powder C- H900 has weaker global texture compared to Wrought-H900 sample by disappearing the BCC ferrite peaks at (200), (211) and (220) along with the reduced relative intensity of (110) peak (Fig.

6. XRD patterns of all three Ar atomized SLM processed and wrought samples in the as- built/as-received and H900 conditions404248Table 2. Measured austenite volume fraction from XRD and EBSDXRDEBSDAs-built H900 As-built H900Powder ANDND<1Powder B~2.6ND<1Powder CNDND~3.7~2.3WroughtNDND~1.2*- not measured, ND- not detected47250Mechanical PropertiesTypical engineering stress-strain curves for SLM and wrought samples in both the as-built and H900 conditions are represented in Fig. 7(a, b). Among the three SLM samples, Powder C-AB sample has higher yield strength with poor elongation and non-uniform plastic deformation in the as-built condition. Overall, the shape of the stress-strain curves for Powder C-AB and Wrought- AR samples are similar that exhibiting a non-uniform plastic deformation from untempered martensitic phase constituent. All the measured mechanical properties are detailed in Table 3.

All the measured mechanical properties are detailed in Table 3. H900 increased the yield strength for all the samples including wrought. Total elongation for all H900 SLM samples weakened after uniform plastic deformation. This observation is consistent with SLM processed 17-4 PH [4]. This could be due to the internal stresses and porosity from the layer-by-layer fabrication of the SLM process. No strain hardening behavior is observed in any of these samples. This can be attributed to the insignificant amount of retained or reverted austenite, which is not sufficient to induce any transformation-induced plasticity (TRIP). Wrought-H900 sample exhibited higher elongation, approximately 17.3%, whereas SLM fabricated sample did not exhibit any improvement in the elongation from H900.

Wrought-H900 sample exhibited higher elongation, approximately 17.3%, whereas SLM fabricated sample did not exhibit any improvement in the elongation from H900. The ultimate tensile strength of Powder C-H900 (~1467 MPa) is comparable to Wrought-H900 ((~1523 MPa) and is higher than the material standard minimum (~1310 MPa) for H900 followed by a solution heat treatment [20]. Tensile and yield strengths of Powder C-H900 are higher than many reported tensile and yield strengths of SLM processed solutionized and hardened 17-4 PH [4, 13, 21, 22]. Despite lower yield and tensile strengths for Powder A and Powder B samples, the total elongation was observed to be higher in the as-built and H900 conditions compared to Powder C samples. Unlike wrought samples, SLM samples total elongation to failure is not improved with H900 from the as-built condition. However, direct hardening using H900 condition enhanced the uniform elongation for Powder C sample.

7. Representative tensile engineering stress-strain curves for Powder A, Powder B, Powder C and wrought samples (a) as-built, (b) H900474Table 3. Summary of mechanical properties showing an average of three measurementsSampleYield strength, 0.2% elastic offset (MPa)Ultimate tensile strength (MPa)Total elongation (%)Micro-Vickers hardness (HV)Powder A-AB607 ± 14763 ± 1422.3 ± 0.7277.3 ± 10Powder B-AB699 ± 17871 ± 2519.4 ± 1.0330.7 ± 9Powder C-AB723 ± 13917 ± 2710.9 ± 0.9333.0 ± 5Powder A-H900873 ± 7953 ± 2416.8 ± 5.4356.1 ± 10Powder B-H9001121 ± 151196 ± 2012.3 ± 2.8396.3 ± 9Powder C-H9001308 ± 301467 ± 119.3 ± 1.6461.9 ± 5Wrought-AR841 ± 131080 ± 1313.1 ± 1.1353.9 ± 9Wrought-H9001371 ± 521523 ± 1817.3 ± 1.2480.2 ± 13ConclusionsIn this study, microstructures and mechanical properties of SLM processed 17-4 PH were evaluated with respect to initial powder chemical composition.

Mechanical performance varied greatly as a result of the evolution of phases resulting from the disparity in the powder chemical composition and H900 heat treatment. H900 hardening - without prior solutionizing - of SLM sample with low Creq/Nieq powder resulted in yield and tensile strengths, and microhardness values that are comparable to wrought samples in H900 condition. Wrought total elongation was greater than that for all SLM samples under all conditions. Regardless of the powder chemical composition, no austenite reversion was observed after H900 treatment.AcknowledgmentsThe U.S Naval Air Systems Command (NAVAIR) under contract number N68335-18-C-0020 funded the research work reported in this paper. We would like to acknowledge Mitch Gennocro from NAVAIR for serving as the technical mentor for the STTR project. We are grateful to Dr. Keng Hsu and Dr. Anagh Deshpande for their initial support in EBSD characterization.References475[1] S. Cheruvathur, E.A. Lass, C.E.

fect of Powder Compaction in Plastic Laser Sintering FabricationToshiki NIINO and Kazuki SATOInstitute of Industrial Science, The University of Tokyo, Japan4-6-1 Komaba Meguro Tokyo, 153-8505 Japanniino@iis.u-tokyo.ac.jpABSTRACT Powder compaction is introduced into plastic laser sintering fabrication. Compaction wascarried out by using a roller of which rotation speed is independently controlled of itstraversing speed. This additional process improved packing density of powder bed by afactor of 20% and reduced residual porosity of obtained parts by a factor of 30%. As anadvantage, powder compaction canimprove mechanical strength of parts ofsemi-crystalline powder, but increases excessive sinter to reduce fabrication accuracyespecially in fabrication of amorphous plastic.

This paper presents characteristics of thepowder compaction process itself and its effects on performance of obtained parts.INTRODUCTION Laser sintering freeform fabrication (“LS” in the following discription) is a relatively newtechnology among those processes that produce three dimensional objects from powderymaterial[1]. Although most of the other powder based production methods includecompaction process before sintering, powder in LS is not explicitly compressed. There aremany reasons for this difference. Difficulty in adding such function to LS systems due tocomplexity of mechanism and high temperature in their process chambers to fulfill the lasersintering process are two reasons for the absense of a compacting process. On the otherhand, less neccessity of compacting in LS than the other powder sintering processes, in which powder must be compressed by high pressure to form green parts, is another reason.

On the otherhand, less neccessity of compacting in LS than the other powder sintering processes, in which powder must be compressed by high pressure to form green parts, is another reason. However, the advantage of powder compaction in LS has not yet denied. In this paper, an experimental method that enables powder compaction without adding major change in mechanism to commercially available system is introduced. The method is tested and improvement in density of the powder bed is measured. Porosity and mechanical properties of processed sinters are evaluated.

The method is tested and improvement in density of the powder bed is measured. Porosity and mechanical properties of processed sinters are evaluated. Effect of different powder material characteristics is also discussed; Both of semicrystaline plastic, PA12 (DuraForm® PA), and amorphous plastic, Polystyrene (CastForm® PS), are tested, and difference in effect of compaction is discussed.POWDER COMPACTIONCOMPACTION METHOD AND DEFINISIONS Compare to normal powder press applied to form green parts using a mold the cross sectionof powder bed in LS system is much larger, thus it is difficult or almost impossible toprovide the powder bed with a high pressure (a large pressing force) at one time with apiston or a flat plate. Instead of such devices, a roller, which is utilized to spread powder byits counter-rotating motion[2] (Fig. 1(a)) can be applied in some LS systems.

Instead of such devices, a roller, which is utilized to spread powder byits counter-rotating motion[2] (Fig. 1(a)) can be applied in some LS systems. To compactpowder bed, we should not rotate the roller in the counter direction as we often find but inthe right direction as shown in Fig. 1 (b). This roller operation introduces powder below theroller to increase the amount of powder under the roller in comparison with the oppositerotational direction’s case, and applies presseure to the powder bed resultantly. Fig. 2 showsresult of compaction with right-direction-rotating roller. As shown here, craters occur on thepowder surface. These craters are created since the powder is compressed so much that thecompressed lump of the powder sticks on the roller surface. This problem can be solved bylimitting the amount of powder in front of the roller. More specifically, recoating processshould follow the scenario shown in Fig.3.

This problem can be solved bylimitting the amount of powder in front of the roller. More specifically, recoating processshould follow the scenario shown in Fig.3. For the first, part piston is lowered by a depth(a) Counter Rotation(b) Rotation in Right DirectionFig. 1 Two types of roller operations(a) Roughen surface of powder bed(b) Powder stuck on the rollerFig. 2 Recoating with roller rotating in right directionFig. 3 Compacting Method(supply thickness, ds, in the following description) that is larger than layer thickness, dl. Then, counter-rotating roller spread the powder and fill the gap created by part piston’slowering in the previous procedure. Reversely, the piston is raised up to the level that islower than the initial state by the layer thickness. Finally, the roller traverses the powderbed to compress the powder as rotating in the right direction.

Finally, the roller traverses the powderbed to compress the powder as rotating in the right direction. Here, we difine “Compaction Factor,” fc, as(cid:1858)(cid:3030) (cid:1568)(cid:3031)(cid:3294) (cid:3031)(cid:3287)(cid:3398) 1 ········································ ①Compaction factor indicates how much the powder bed is compacted, and it becomes zerowhen supply depth and layer thickness becomes the same.PACKING DENSITY OF UNSINTERED POWDER To evaluate improvement in packing density of powder by the compaction method,experiments using an LS system were carried out. A hollow box as shown in Fig.

A hollow box as shown in Fig. 4 is builtand packing density, c, of the unsintered powder (cake in the following description) remaining in the cavity is calculated by following fomula.(cid:2025)(cid:3030) (cid:3404)(cid:3040)(cid:3281)(cid:2879)(cid:3040)(cid:3280) (cid:3023)(cid:3278) (cid:3096)(cid:3295)········································ ②where mf, me, Vc and t are mass of the built box including the cake, mass of the emptied box, volume of the cavity calculated from measurement of its inner dimensions and specificgravity of the powder material, respectively.Fig. 5 shows relationship between, packing density and compression factor. In theseFig. 4 Test specimen for packing density measurement50]%[48e k a C46f oy t i s n e D g n i k c a P4442403800.5 1.5 1.0 Compaction Factor2.0Fig. 5 Relationship between Packing Density and Compaction FactorTbl.

5 Relationship between Packing Density and Compaction FactorTbl. I Characteristics of PA powderAverage grain size58μm*Apparent Density (ISO 60)0.42 g/cm3 (40%)Tapped Density0.50 g/cm3 (50%)Specific gravity1.00 g/cm3*Data provided by 3D systemsexperiments, a polyamide powder (DuraForm® PA from 3D Systems) is used. Tbl. I summarises characteristics of the powder. A prototype version of Semplice (ASPECT Inc.),which is equipped with a roller system that can control rotation speed and travesing speedof the roller independently, was employed. Layer thickness of 100m was selected. Traversing speed, vct, and rim speed, vcr, of the roller during compacting process share the same value of 100mm/s to avoid friction between the roller and powder bed. Packing density increased with compaction factor and reached the highest value of 48% when fc = 2.0. This value is larger than in the case that compaction is not applied but still lower thantapping density of 50%.

This value is larger than in the case that compaction is not applied but still lower thantapping density of 50%. Higher compaction factor causes failure such as crater and drag ofthe part bed.Fig. 6 shows relationship between packing density and translating speed of the roller invarious compaction factors. The lower roller speed is, the higher packing density becomes. We can find that the highest density is 49% when vct=50mm/s and fc = 2.0. However, this compaction condition cannot be used since it causes drag of the previous layer at a certainprovability.ROTATION SPEED OF ROLLER IN COMPACTION PROCEEDURE In experiments of the previous section, rotation speed of the roller was controlled so that itsrim speed becomes the same.

Following is discussion on phenomena that occur when thetwo speed do not match.52Compaction factors]5048=2.0 f c =1.0 f c Not Applied%[y t i s n e d g n i k c a P4644424038050100150200250Translating speed at compacting [mm/s]Fig. 6 Relationship between packing density and transverse speedFig. 7 includes side views of sinters obtained in various conditions of rotation speed andtraversing speed. Translating speed of 50mm/s was used for photos (b) through (e). Rimspeed was varied. Compaction factor is 2.0 for all photos. The LS system employed in thisresearch has a pair of feed cylinder on the both sides of its part cylinder. In normal processwithout applying compaction process, the both cylinders are used alternately. Contrary,when compaction process is added, the roller go from one side to the other to spread thepowder, and returns to compact the bed.

Contrary,when compaction process is added, the roller go from one side to the other to spread thepowder, and returns to compact the bed. Thus, only one side of the feed cylinders is used ifwe do not switch the spreading and compacting direction intentionally. Since capacity ofa feed cylinder of the employed machine in this research is insufficient to complete one testpiece, the feed cylinder is switched during a build. Bold lines in the pictures in Fig. 7display layer where the directions were switched, and arrows indicate the compactingdirections. Followings are explanations for results sown in Fig. 7. (a) fc = 2.0, vct= vcr=100mm/sRefference for successful build.

Followings are explanations for results sown in Fig. 7. (a) fc = 2.0, vct= vcr=100mm/sRefference for successful build. (b) fc = 2.0, vct=50mm/s, vcr=25mm/sPreviously built object is dragged in theforwad direction.(a) fc = 2.0, vct= vcr=100mm/s(c) fc = 2.0, vct=50mm/s, vcr=0mm/sDrag in the same direction is observed but ismilder since not rotating roller does notintroduce so much powder.(b) fc=2.0, vct=50mm/s, vcr=25mm/s(d) fc = 2.0, vct=vcr=50mm/sAlthough the rubbing between the roller andbed is minimized, small drag is still observed.We cannot find the relation between the direction of the drag and roller motion. As a result of compressing large amount of powder under the roller, pressure rose to very high level, and a friction force drags the previous layer in an unpredictable direction.

As a result of compressing large amount of powder under the roller, pressure rose to very high level, and a friction force drags the previous layer in an unpredictable direction. (e) fc = 2.0, vct=50mm/s, vcr=100mm/s(c) fc=2.0, vct=50mm/s, vcr=0mm/s(d) fc=2.0, vct=vcr=50mm/sSince rim speed is faster than traversing speed,(e) fc=2.0, vct=50mm/s, vcr=100mm/sdrag in the backward direction occurred.Fig.7 Drags caused in various roller speed conditionsFrom these results, we obtained the following conclusions. 1. We should equalize the roller’s rim speed with its translating speed to avoid drag. 2. Insufficient and excess speed of the roller results in the drag of the lower layers in frontand back, respectively.3. High packing density raises the provability of drag even if rim speed is in accordancewith traversing speed.EFFECT OF POWDER COMPACTIONIMPROVEMENT IN PACKING DENSITY OF SINTER Fig.

High packing density raises the provability of drag even if rim speed is in accordancewith traversing speed.EFFECT OF POWDER COMPACTIONIMPROVEMENT IN PACKING DENSITY OF SINTER Fig. 8 shows relationship between packing density and compaction factor at various laserpowers. The packing density was obtained by meausrement of mass and dimensions of test specimen. Packing density increses with compaction factor unitil fc > 1.0. When fc > 1.0, the density decreases in turn since amount of excessive sinter became large. The maximumpacking density is 94% while one at the same building condition without compaction is91%. Though improvement of 3% in packing density seems quite small, relative reductionrate in porosity of 33% is not a small value.REDUCTION OF CURL DISTORTION In this section, we discuss the effect of powder compaction on occurrence of curl distortion. Cubic specimens with dimensions of 4mm×10mm×80mm were built from PA12 powder.

Cubic specimens with dimensions of 4mm×10mm×80mm were built from PA12 powder. Heat barriers, thin disks which are built below objective sinter for the purporse ofpreheating, were not used to obtain measurable distortion. Photos for the specimens built inthe both conditions are shown in Fig. 9. Curl distortions were successfully suppressed, andthe measured value when compacting is applied and not applied are 0.3mm and 0.8mm,respectively. We can roughly categorize curl distortion into two groups. The first one includes curls that are caused during repeating powder supply and laser exposure. We name this type of curls as “curls in process”. The other one includes those that occur during cooling stage of the powder bed, after layering is finished. We call this category as “curls in cooling.” Since curls in cooling occur after objects are build correctly, thickness is constant over all the area when a flat plate is built as shown in Fig. 10(a).

We call this category as “curls in cooling.” Since curls in cooling occur after objects are build correctly, thickness is constant over all the area when a flat plate is built as shown in Fig. 10(a). On the other hand, the plate is thinner near the edge in the case of curls in process as shown in Fig. 10(b), since the curls of the layer occurred earlier, i.e. lower layer, is larger. In LS process, solidification of melted orpartially melted powder on the objects being solidified previously, repeats in layer by layerfashion. Since solidification involves shrinkage, LS includes process to generate curldistortion inherently. To avoid this dilemma, various countermeasures are taken incommertially available systems and materials. Among these measures, controlling reducethe curls. LS system’s recent trend toward scalling-up of build-envelope raises difficulty ofprecise temperature control over the whole build area.

Among these measures, controlling reducethe curls. LS system’s recent trend toward scalling-up of build-envelope raises difficulty ofprecise temperature control over the whole build area. If powder compacting can reduce the1009510W 15W 17W]%[y t i s n e D g n i k c a P9085807500.51.0Compaction Factor1.52.0Fig. 8 Relationship between packing density of sinter and compaction factora. Top viewb. Side viewFig. 9 Photos of specimens for curl distortion measurement. Top: without compacting, Bottom: with compacting.tetctetca. Curl in cooling (tc  te)b. Curl in process (tc > te)Fig. 10 Two types of curl distortioncurl distortion when powder bed temperature is out of the range, application of compactingis helpful for development of today’s large scaled machines. To evaluate this effect, cubes with dimension of 52mm×52mm×11.5mm were built at various temperature. The results are listed in Tbl. II. Tbl.

To evaluate this effect, cubes with dimension of 52mm×52mm×11.5mm were built at various temperature. The results are listed in Tbl. II. Tbl. III shows curls of the specimen built at 165C. As shown here,powder compaction successfully reduced the distortion that was caused by insufficientpowder bed temperature. Though residual curl of 0.2mm when compaction is applied is notacceptable yet, this can be suppressed by taking other supplementary countermeasures suchas heat barriers. This result demonstrates that powder compaction can expand the width ofthe process window from 5K to 10K.IMPROVEMENT IN MECHANICAL PROPERTIES Effect of powder compaction on improvement in mechanical properties was investigated.Tbl. II Building Result in Various Temperature ConditionsPowder bed temperature [C]Result180Powder bed melted.175Good170Good165Building is finished, but curl occurred.160Collision between roller and sinter. Building was failedTbl.

Building was failedTbl. III Reduction of curl in building at 165CCompaction FactorCake Density [g/cm3]Sinter Density [g/cm3]Height at the center [mm]Height at the edge [mm]Curl [mm]No0.400.9111.310.60.7Compaction2.00.460.9211.511.30.2Tbl. IV Improvement in Mechanical Strengthw/o Compaction with compactionincrease Rate of increaseYield Stress [MPa]42.044.02.05%Elongation @ break [%]13.814.40.64%Flexural stress [MPa]51.156.15.010%Flexural Modulus [MPa]1178134716914%Impact strength (not notched) [kJ/m2]2736933%Tensile test (ISO 527), three-point bending test (ISO 178) and Izod impact test (ISO 180)were performed (Tbl. IV). Impact strength improvement is the most significant, and otherssuch as bending are following. Performances against tensile stress are also improved, butthe extent of improvement is much smaller.

Impact strength improvement is the most significant, and otherssuch as bending are following. Performances against tensile stress are also improved, butthe extent of improvement is much smaller. In comparison with a small improvement indensity of 1% the strength is relatively large although a strong relationship betweenpacking density and mechanical properties of a powder [3].

In comparison with a small improvement indensity of 1% the strength is relatively large although a strong relationship betweenpacking density and mechanical properties of a powder [3]. Here, we can guess thatcompressing force during compacting process increased adhesion between grains or layers.In addition, better improvement in strength against bending than against pulling indicatesthat improvement of the adhesion between layers occurred, and it supports our hypothesisthat pressing force is major cause of strength improvement.COMPACTION OF AMORPHOUS PLASTIC Although LS process is applicable to all thermoplastics in principle, choices of plastics thatcan be used for fabrication of structural parts are quite limited to some semicrystalinepolymers such as polyamide and polypropylene.

In this section, we discuss the applicationof powder compaction to LS process of amorphous plastic which is mostly used forfabrication of lost models for casting.IMPROVEMENT IN PACKING DENSITY OF POWDERPolystyrene powder (CastForm™ PS, 3D Systems) is used in the following tests. Tbl. V shows characteristics of the powder compared to polyamide powder discussed in theprevious section. Tbl. VI shows the improvement of density of the powder bed. For calculation of packing density, specific gravity of 1.05g/cm3 is used as typical value for polystyrene. A compaction factor of 1.0 can raise the packing density from 39% to 44% (0.46g/cm3 in density). This value is equivalent to tapped density. Compaction at higher compaction factor was not successful. These experimental results show that PS powder can be compacted more easily than PA powder.

This value is equivalent to tapped density. Compaction at higher compaction factor was not successful. These experimental results show that PS powder can be compacted more easily than PA powder. We can not specify where this difference derived from; nature of the materials or acquired characteristics when the material is shaped into powder grains.IMPROVEMENT IN DENSITY OF SINTER AND EXCESSIVE SINTER Density of processed sinter is measured. As a test specimen, a cube with dimensions of 52mm×52mm×11.5mm is used. Tbl. VII shows packing density and dimensions ofTbl. V Characteristics of PS powderPSPAAverage grain size62μm*58μm*Apparent Density0.40 g/cm30.40 g/cm3Tapped Density0.46 g/cm3*0.50 g/cm3Data provided by 3D SystemsTbl. VI Packing Density of PS powder after compactionPacking Density of PowderCompaction FactorPSPANot applied39%41%1.044%(46g/cm3)44%2.0‐46%Tbl.

VI Packing Density of PS powder after compactionPacking Density of PowderCompaction FactorPSPANot applied39%41%1.044%(46g/cm3)44%2.0‐46%Tbl. VII Packing density and dimensions of sinter of PS powderCompactionWidth in xaxisWidth in yaxisWidth in zaxisPackingdensityNot Applied51.6mm51.6mm11.9mm53%fc = 1.051.8mm51.6mm12.1mm56%processed sinters. Compaction of fc = 1.0 improved density of sintered by 3% to 56%. This packing density is still so low that we cannot use these sinters as structural parts. Inaddition, compacting adds excessive sinter reducing accuracy of the parts.We can conclude that compaction does not bring positive change to process withamorphous plastic.DISCUSSIONS Though, in this research, counter-rotating roller is employed for powder spreading, we can replace this mechanism with a simple blade and place it just before the compacting roller.

Then, powder spreading and compaction is performed simultaneously in a single roller traversing action. This will be able to halve the recoating time, consequently. Compacting phenomena having been discussed in this paper requires elasticity of the powder or roller mechanism. Otherwise, compaction factor must be equal to increase rateof powder density. Fig. 11 explains the case that elasticity of powder bed plays dominantFig. 11 Schematic view for compaction at high compaction factorrole in solving the problem. Surface level of powder bed before compacting roller is veryhigh. When the roller runs over the high powder bed, the powder is once compressed sothat the level is lowered to the bottom line of roller. In this moment, a large amount ofpowder exists under the roller, but the stress is relieved by deformation of the powder in thewide range around the roller.

In this moment, a large amount ofpowder exists under the roller, but the stress is relieved by deformation of the powder in thewide range around the roller. The deformation derives from plasticity and elasticity of thebed, and deformation owing to the latter one is released and powder level is partiallyrecovered after the roller passed.For analysis of compacting powder, we have to establish new analytic model in the future.Shanjani et al introduced a mathematical model [2], but the model is based oncounter-rotating roller system though we need one for the roller rotating in the reversedirection. We find similar mechanism in powder rolling process and analytic model isprovided [4, 5]. In this process, powder is compressed to form a seat as it is sandwiched andconveyed by a pair of rotating rollers. It seems possible to use this model for ourmechanisms.

In this process, powder is compressed to form a seat as it is sandwiched andconveyed by a pair of rotating rollers. It seems possible to use this model for ourmechanisms. However, modification is still required since phenomenon is slightly different.For example, thickness of the seat is very much larger in our system. Speeds at whichpowder is involved and seat is rolled out are the same.Quite interesting is that density of the unsintered powder bed is lower than apparent density when compacting is not performed. This means counter-rotating roller spreads and lays the powder quite loosely. This is quite adequate to recoating new powder in very thin layer uniformly without giving drag force to sinters in the previous layers. However, it is not good for adhesion since adding pressure is an essential prerequisite for adhesion. In that meaning, adding pressure after forming layer by counter-rotating roller is quite reasonable for the both of quality of layer and adhesion.

In that meaning, adding pressure after forming layer by counter-rotating roller is quite reasonable for the both of quality of layer and adhesion. Compaction is quite effective on improvement of adhesion but increases excessive sinter as its drawback. Since these merits and demerits derive from the same phenomenon, we haveto make some trading-off to use compaction process. It is quite dependent on thecharacteristics of material, and in the case of amorphous plastic, we should not apply thenew supplemental process.CONCLUSION An experimental compaction method was introduced. The method consists of two processes.In the first process, a counter-rotating roller spreads the powder, and in the later process, thepart piston is heightened and rotor rotating in the right direction compresses the powdersurface. This compaction process successfully raised powder density of the bed to itstapped density.

This compaction process successfully raised powder density of the bed to itstapped density. To avoid drag of the powder bed, rim speed and traversing speed of theroller in compaction process should be equalized, and powder bed density should be lowerthan tapped density. In case of polyamide, powder compaction reduces the curl distortionand improves mechanical properties up to 33%. Reduction of curl distortion can double thewidth of process window of part bed temperature. Improvement of mechanical propertyderives from that in adhesion by adding pressure to the powder. Application of compactionto amorphous powder increases excessive sinter and we could not find its merit.REFERENCES [1] Beaman J, et al, “Solid Freeform Fabrication: A New Direction in Manufacturing,”1996[2] Shanjani Y, et al, “Material Spreading and Compaction in Powder-Based Solid FreeformFabrication,” Proc.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of Powder Degradation on the Fatigue Behavior of Additively ManufacturedAs-Built Ti-6Al-4VPatricio E. Carrion1, 2, Arash Soltani-Tehrani1, 2, Scott M. Thompson1, 2, Nima Shamsaei1, 2, * 1National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849, USA 2Department of Mechanical Engineering, Auburn University, Auburn, AL 36849, USACorresponding author: shamsaei@auburn.eduAbstractAdditive manufacturing (AM) technology has enabled many industries to generate functional parts with an increased level of complexity via a layer-by-layer melting. In laser-powder bed fusion (L- PBF), the most commonly used AM process for metals, powder is often recycled due to its high cost.

In laser-powder bed fusion (L- PBF), the most commonly used AM process for metals, powder is often recycled due to its high cost. However, there is no comprehensive study on how recycling powder affects its rheological properties, and the mechanical and fatigue behavior of the final manufactured part. In this study, a comparison of new and used Ti-6Al-4V powder characteristics was made. The comparison includes morphology, size distribution, as well as monotonic tensile and fatigue behavior of fabricated specimens. Conclusions and suggestions on powder recycling are made. Results indicate that the powder particle size distribution (PSD) becomes narrower and the morphology of the particles change with recycling.

Conclusions and suggestions on powder recycling are made. Results indicate that the powder particle size distribution (PSD) becomes narrower and the morphology of the particles change with recycling. However, no comparable effect was observed on the monotonic tensile and fatigue behavior of the AM as-built Ti-6Al-4V specimens.KEYWORDS: Laser-Powder Bed Fusion, Ti-6Al-4V, Fatigue Life, Powder Degradation, Powder Recycling.IntroductionAdditive manufacturing (AM) has provided the opportunity to generate complex parts in an integrated unit. This has helped eliminate the need for processes to join parts including brazing and welding. The two most common AM methods for metals are direct energy deposition (DED) and powder bed fusion (PBF). In the PBF process, unlike subtractive conventional manufacturing, parts are fabricated in a layer-wise system via metallic powders. Laser powder bed fusion (L-PBF), a type of PBF process, uses a laser as the energy source.

Laser powder bed fusion (L-PBF), a type of PBF process, uses a laser as the energy source. In this process, the metal powder is spread over the substrate, and the laser initiates a continuous, selective fusing process. After the desired part is fabricated, the remaining powder can either be discarded or recycled. If recycled, the powder from the collector bin is sieved to remove spatter or debris, and then mixed with any remaining or unused powder in the feed bin. Due to the high cost of metallic powders, specifically Ti-6Al-4V alloy, powder recycling can be a viable option for cost reduction. Few studies have been conducted on the effects of powder recycling on final mechanical properties, and there is no specific procedure on how to recycle nor quality control.1366Powder is defined as a finely divided solid which is smaller than 1 mm in its maximum dimension [1]. A particle is the smallest unit of the powder which cannot be subdivided.

A particle is the smallest unit of the powder which cannot be subdivided. Powder properties such as size distribution, flowability, and chemical content have been shown to affect the mechanical properties of the AM fabricated parts [2–6]. For instance, it has been observed that a 0.15% water amount added to a powder can result in agglomeration in metallic powders and consequently powder flow is affected [1].Jacob et al. [2] studied the mechanical properties, including tensile and hardness, on 11 similar L-PBF builds using nitrogen gas atomized 17-4 PH stainless steel. It was observed that the recycled powder had a better flow time, increased powder bed density and apparent density. However, no comparable change was observed in powder size distribution (PSD), and shape. Also, the martensitic-ferritic phase was slightly increased due to fabrication and recycling in the nitrogen atmosphere.

Also, the martensitic-ferritic phase was slightly increased due to fabrication and recycling in the nitrogen atmosphere. The parts fabricated from recycled powder showed the same mechanical properties as the parts from new powder.Liu et al. [3] studied two stainless steel 316L batches from LPW Technology Ltd and Sandvik Osprey Ltd with PSDs of, respectively, 15-45 µm and 0-45 µm, to study the effects of PSD on the processing parameters optimization in L-PBF and final part properties. It was observed that powders with narrower range of PSD have better flowability – leading to parts with higher ultimate tensile strength (UTS) and increased hardness. It was seen that a wider PSD leads to higher densities since finer particles can fill in the vacant areas between larger particles. It was also found that the powder with a wider PSD had a smoother surface finish. However, wider PSDs were shown to adversely affect agglomeration due to the higher friction between particles.Seyda et al.

However, wider PSDs were shown to adversely affect agglomeration due to the higher friction between particles.Seyda et al. [4] investigated the effects of aging on Ti-6Al-4V powder. Part properties including density, porosity, hardness, and surface quality were studied to understand the effects of aging. It was found that the particles coarsened and flowability increased due to the apparent density increase and PSD enlargement after recycling multiple times. No significant effect was seen on the porosity, density, surface roughness, hardness and mechanical strength for the aged manufactured specimens. However, less pore locations were found in the bulk material of the aged powder with increased pore sizes. Also, there was an increasing trend in the surface roughness values with aging.O’Leary et al. [5] observed that the surfaces of particles roughened after recycling Ti-6Al- 4V powder five times.

Also, there was an increasing trend in the surface roughness values with aging.O’Leary et al. [5] observed that the surfaces of particles roughened after recycling Ti-6Al- 4V powder five times. Also, it was seen that the number of finer particles (< 15 µm) decreased while larger particles (> 45 µm) increased. No significant change in amount of oxygen and nitrogen was seen. However, a comparable amount of oxygen absorption was observed in the parts while the fabrication atmosphere was argon.Tang et al. [6] studied the effects of powder recycling on the Ti-6Al-4V powder quality and parts fabricated via selective electron beam melting (SEBM). It was observed that the oxygen content slightly increased after recycling 21 times while the Al and V contents were almost constant with a negligible decrease. It was also seen that the powder particles became less spherical.

It was also seen that the powder particles became less spherical. No significant effect was observed on tensile properties of produced specimens.The aforementioned studies have demonstrated changes in powder characteristics due to recycling and the subsequent effects on the mechanical properties of fabricated parts. However, these parts are often subjected to dynamic loading; therefore, there is a further need to evaluate the fatigue performance of parts fabricated using recycled (used) powder. By incorporating unique powder/part characterization methods, this study quantifies the rheological properties of1367continually recycled Ti-6Al-4V powder while focusing on the PSD and shape.

By incorporating unique powder/part characterization methods, this study quantifies the rheological properties of1367continually recycled Ti-6Al-4V powder while focusing on the PSD and shape. The AM specimens are subjected to quasi-static tensile and fatigue testing, the results are then correlated to the changing powder quality.Experimental ProgramPowder Characterization The material selected for this study was gas-atomized Ti-6Al-4V Grade 23 powder with a particle size of 15-45 µm, manufactured by LPW Technology. The powder was in two conditions, new (i.e. virgin) and used (recycled more than 14 times). It is important to mention that the recycling procedure involved sieving the used powder with an 80-µm mesh size sieve after each print, and if needed, new powder was added in order to have a sufficient amount to complete the next part/component fabrication.

Particle size distribution (PSD) analysis was conducted via a Malvern Mastersizer 3000 Hydro EV laser diffraction particle size analyzer and based on ASTM E2651-13 [7]. The powder used for size analysis was sampled according to ASTM B215-15 by taking equal increments from different depths of the container [8].Material and Specimen Fabrication L-PBF Ti6Al-4V specimens were fabricated using an EOS M290 system in two separate batches, one with new and one with used powder. EOS default process parameters were used, including 280 W laser power, 1200 mm/s travel speed, 100 mm stripe width, 140 µm hatch spacing, 67° hatch rotation, and a 30 µm layer thickness. The specimens were fabricated vertically (90° from the build plate) with argon as the shielding gas. The specimen geometry was based on ASTM E606 standard [9], which includes a uniform gage section with a 3.75 mm gage diameter, and a 15 mm gage length.

The specimen geometry was based on ASTM E606 standard [9], which includes a uniform gage section with a 3.75 mm gage diameter, and a 15 mm gage length. After fabrication, specimens were annealed at 704 °C in an argon atmosphere for one hour, and then air cooled to the room temperature to remove residual stresses. Specimens were then tested without further post-processing and considered to be in the “as-built, annealed” condition. It is important to emphasize that the only variable was the Ti-6Al-4V powder (i.e. new and used), as great effort was employed to keep all other variables such as specimen layout on the build plate, shielding gas, etc., consistent; to only investigate the powder recycling effects on the mechanical properties and fatigue performance.Testing Procedures Strain-controlled quasi-static tensile and fatigue tests were conducted using an MTS 810 servohydraulic system, and an MTS extensometer with a 10 mm gage length was used to record strain.

Tensile tests were performed at a 0.001 s-1 strain rate on two specimens fabricated with the new powder, and two with the used powder. Fully-reversed (Rε = -1) strain-controlled fatigue tests were conducted according to ASTM E606 standard [9]. The strain amplitudes values selected for this study were 0.001, 0.003, and 0.005 mm/mm. The test frequency was adjusted for each strain amplitude to eliminate strain rate effects on the fatigue behavior. Experiments that reached over 4x106 reversals were determined to be a run-out. Three fatigue tests for each strain amplitude were conducted for verification of results. Specimens from the same location on the build plate, but from different powder batches (new and used), were tested with same parameters.1368Experimental Results and AnalysisPowder Recycling Comparison The particle size distribution (PSD) of both powder samples are presented in Fig. 1. It may be seen that there is a difference in PSD between the new and used powders.

1. It may be seen that there is a difference in PSD between the new and used powders. The size distribution becomes narrower for the used powder. This is explained by the gap between the wiper/recoater and the substrate being smaller than 50 µm, hence finer particles are spread on the build plate and participate in the melting and solidification processes during manufacture. However, particles greater than 50 µm will be transferred to the collector bin. After completion of each print, the powder is sieved and mixed with the remainder of the powder in the feed bin. Repeating this procedure will lead to narrower PSDs because finer particles are melted and solidified, and larger particles get detached when passing through the 80 µm mesh size screen. The average diameter of used powder particles was slightly less than the new powder particles.)%(y t i s n e D e m u l o V18 16 14 12 10 8 6 4 2 0New Used100 Particle Size(µm) Fig. 1.

The average diameter of used powder particles was slightly less than the new powder particles.)%(y t i s n e D e m u l o V18 16 14 12 10 8 6 4 2 0New Used100 Particle Size(µm) Fig. 1. Particle size distribution (PSD) comparison of the new and used Ti-6Al-4V powders.1101000Representative powder morphology, obtained via scanning electron microscopy (SEM), is shown in Fig. 2. Results indicate that fewer “satellites”, smaller particles attached to the larger particles, were observed in the used powder when compared to the new one. These particles may contribute to defect or pore formation since they leave void spaces by spattering from the deposited layer when the laser scans the selected area. It was also observed that most particles are still spherical after recycling the powder; however, similar irregular shape particles were observed in both new and used powder.(a)(b)Fig. 2.

It was also observed that most particles are still spherical after recycling the powder; however, similar irregular shape particles were observed in both new and used powder.(a)(b)Fig. 2. Particle morphologies obtained via SEM for (a) new and (b) used Ti-6Al-4V powder.1369Tensile Deformation Quasi-static tensile stress-strain curves of L-PBF Ti-6Al-4V specimens fabricated with new and used powders are presented and compared to wrought Ti-6Al-4V [10] in Fig. 3. It may be seen that the yield, and ultimate tensile strength of L-PBF Ti-6Al-4V are higher (10%, and 5% respectively) than the wrought counterpart, but the elongation to failure is comparable. It is also shown that the ultimate tensile strength of the new powder specimens is 5% higher than the used powder, while the yield strength and elongation to failure differences were negligible. These results differ from other studies such as Tang et al.

These results differ from other studies such as Tang et al. [6], where it was determined that for Ti-6Al- 4V parts fabricated via SEBM, the yield and ultimate tensile strength increased with increased powder usage due to higher oxygen content. These results were also determined to be of negligible impact on the overall mechanical properties. Jacob et al. [2] however, presented findings for 17-4 PH stainless steel, where it was concluded that the effects of powder recycling on the mechanical properties are negligible. From the quasi-static tensile experiments, it can be concluded that effects of powder recycling do not have significant effects on the mechanical properties of L-PBF Ti-6Al- 4V.1,2001,000) a P M800(600 σ, s s e r t S400200New Used Wrought [9][10]0 0.0000.0400.080 Strain, ε (mm/mm)0.1200.160Fig. 3.

3. Tensile deformation comparison of L-PBF specimens fabricated with new and used Ti-6Al-4Vpowder compared to wrought Ti-6Al-4V [10].Fatigue Behavior Figure 4 presents a semi-log plot of the strain amplitude versus reversals to failure of L- PBF Ti-6Al-4V specimens fabricated with new and used powder. It may be seen that there are no significant differences in fatigue life across the three strain amplitudes tested (0.001, 0.003, and 0.005 mm/mm). Horizontal arrows indicate run-outs (i.e. a fatigue life over 4x106 reversals). At a strain amplitude of 0.001 mm/mm, one of the specimens fabricated with the new powder resulted in a run-out while for the used powder there were two. These data points are denoted by the “x3” in in Fig. 4. Minimum differences in fatigue lives could be attributed to the as-built surface roughness condition, which has been found to have significant influence on the fatigue1370performance of AM parts [11].

It provides the opportunity for failure to occur from surface defects which typically have higher stress concentration than internal defects. In conclusion, the fatigue performance of as-built L-PBF Ti-6Al-4V specimens is not affected by powder recycling as substantial differences or scatter in data were not observed.0.006New)m m m m/(0.0050.004x2Useda ε, e d u t i l p m A n i a r t S0.0030.0020.001x2x30.0001.E+31.E+41.E+51.E+61.E+7Reversals to Failure, 2NfFig. 4. Strain-life comparison of L-PBF specimens fabricated with new and used Ti-6Al-4V powder.ConclusionsIn this study, the change in powder particle size distribution due to recycling and its effects on the tensile deformation and fatigue behavior of L-PBF Ti-6Al-4V under strain-controlled fully- reversed (Rε = -1) fatigue loading was investigated. The following is a summary of the main conclusions and novel findings:1.

The following is a summary of the main conclusions and novel findings:1. Fewer satellite/fine particles were observed for the used powder and the particles were near- spherical with fewer irregularly-shaped particles after recycling.2. The particle size distribution became narrower after recycling due to finer particles being melted and solidified while the larger particles are moved to the collector by the wiper.3. Larger particles (>80 µm) available in the new powder get discarded due to the sieving process before each fabrication.4. Powder recycling has negligible effects on the mechanical behavior and fatigue performance of the manufactured as-built L-PBF Ti-6Al-4V specimens. This is explained by the surface roughness having a greater effect in fatigue.1371AcknowledgementsThis research was partially funded by the National Science Foundation (NSF) under grantNo. 1657195References [1] German RM. Powder metallurgy science.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceTHE EFFECT OF POWDER ON COOLING RATE AND MELT POOL LENGTH MEASUREMENTS USING IN SITU THERMOGRAPHIC TECNIQUESJ. C. Heigel* and B. M. Lane*National Institute of Standards and Technology1, Gaithersburg, MD 20899AbstractHigh-speed thermal cameras enable in situ measurement of the temperatures in and around melt pools generated during powder bed fusion processes. These measurements can be used to validate models, to monitor the process, and to understand the microstructure formed during the process. Unfortunately, pre-placed powder layers complicate the measurement due to spatter and irregular surfaces that impact emissivity.

Unfortunately, pre-placed powder layers complicate the measurement due to spatter and irregular surfaces that impact emissivity. The objective of this work is to present high speed thermographic measurements of single and multiple scan tracks on substrates with and without pre-placed powder and to analyze and compare the impacts of powder on melt pool length and cooling rate measurements.IntroductionThermographic measurements of powder bed fusion (PBF) provide valuable insight into the process. The measurements can be used to study the process [1–3], to validate models [4], to better understand the material transformations that occur during the process [5–6], or to develop real-time monitoring and control techniques [7 - 11]. However, optical measurement of the surface temperature is challenging for a variety of reasons, such as the variable emissivity of the surface [12–13] and spatter that regularly occurs during the process [14].

While many studies have implemented thermographic techniques to measure the melt pool length and cooling rates during PBF processes, the impact of the process on the measurement and its error have not been well characterized.Many of the challenges involved with thermography of PBF can be alleviated by removing the powder and studying the process while scanning on bare metal surfaces. Although this strategy removes the experiment from the actual process that consolidates layers of powder, it has two advantages. First, scans on bare surfaces are easier to measure, enabling the measurement itself to be studied.

First, scans on bare surfaces are easier to measure, enabling the measurement itself to be studied. The second advantage is that the simplified process is easier to simulate and allows the modeler to validate the physics and material model without the powder, decreasing the number of variables in the model.For instance, similar to other studies [1, 3, 4, 6, 7, 9–11], the earliest PBF thermography studies performed by the National Institute of Standards and Technology (NIST) were performed during the build of a small part [14–15]. Although informative, the chaotic nature of the spatter and the rough solidified surface made the process of quantifying the measurements difficult. Consequently, the experiments have been simplified by removing the powder so that better quality data could be acquired during single line1 Certain commercial equipment, instruments, or materials are identified in this paper in order to specify the experimental procedure adequately.

Such identification is not intended to imply recommendation or endorsement by the National Institute of Standards and Technology, nor is it intended to imply that the materials or equipment identified are necessarily the best available for the purpose. This publication was prepared by United States Government employees as part of their official duties and is, therefore, a work of the U.S. Government and not subject to copyright.1340Published in the Proceedings of the 2017 Annual International SFF Symposiumscans, enabling the melt pool length to be measured over the processing space [2] and to provide model validation data [5].

Now that the measurement process and thermal behavior of the material are better understood, powder can be re-introduced into the experiment so that its effect on the measurement and the process can be investigated.This objective of this work is to present a preliminary investigation into the impact that powder has on thermographic measurements of melt pool length and cooling rate. Several cases are studied, two single line scans and two multiple scan tracks, half on a bare surface and the other half on a single layer of hand-spread powder. The acquired thermal images are qualitatively compared to understand spatter and the approximate length of the hot region as a function of line count. Radiant temperature profiles are extracted from these images and the cooling rate is calculated.

Radiant temperature profiles are extracted from these images and the cooling rate is calculated. Observations are made from these comparisons and their impact on the current work and on the future direction of the thermographic effort at NIST is discussed.Experimental MethodThe experiments consist of scanning single tracks and small areas using a commercial laser PBF machine and measuring the radiant temperature of the process using a high-speed thermal camera. The camera observes a 12 mm by 6 mm area in the build chamber through a custom fabricated door [2, 14]. The camera operates at a frame rate of 1800 Hz and an integration time of 40 µs, and the system is sensitive to wavelengths between 1350 nm and 1600 nm and can measure radiant temperature between 550 °C and 1038 °C.

Although the camera does not directly measure the true temperature of the surface, the camera signal and surface temperature are related according to [16]:𝑆meas = 𝐹(𝑇radiant) = 𝜀 ∙ 𝐹(𝑇true) + (1 − 𝜀) ∙ 𝐹(𝑇environment)where Smeas is the signal measured by the camera, F(Tradiant) is a calibration function to convert the signal measured by the camera to a black-body temperature, Ttrue is the actual temperature of the target surface, Tenvironment is the temperature of the environment, and ε is the emissivity of the surface. When the emissivity of the surface is unknown, as in this study, the radiant temperature of the surface is reported.Figure 1 presents the setup inside the build chamber. Figure 1A is a view through the viewport on the custom door of the rotary sample holder and five 25 mm square substrates (3 mm thick). Figure 1B shows a close-up of one of the samples after an area has been scanned.Table 1 presents the four different cases studied in this work.

Figure 1B shows a close-up of one of the samples after an area has been scanned.Table 1 presents the four different cases studied in this work. Cases 1 and 2 produce single line scan tracks while Cases 3 and 4 scan a 5 mm x 4 mm area using multiple consecutive scans (39 in total, alternating direction). The scans are offset by 0.1 mm (hatch distance). Cases 1 and 3 are executed on bare substrates whereas Cases 2 and 4 are performed on a single hand-spread layer of powder (approximately 36 µm thick). In each case the nominal laser power is 195 W and the scan speed is 800 mm/s.

In each case the nominal laser power is 195 W and the scan speed is 800 mm/s. In all cases, each scan track is 5 mm long.Table 1 - The test cases investigated in the current work.Number of scan tracks Layer of powder?Case 1 1 NoCase2 1 YesCase 3 39 NoCase 4 39 Yes1341(1)Published in the Proceedings of the 2017 Annual International SFF SymposiumHand-spread powder layerB)A)Five substrates are loaded into holderSubstrate holder mounted to the rotary stage and visible to the thermal cameraScanned area25 mm square substrateFigure 1 - Experiment setup inside the PBF build chamber. A) Image acquired through the viewport on the custom door showing the sample holder and small substrates. B) Close up of a substrate with a hand-spread powder layer after scanning an area (Case 4).ResultsFigure 2 presents select frames acquired during steady-state in Cases 1 and 2.

B) Close up of a substrate with a hand-spread powder layer after scanning an area (Case 4).ResultsFigure 2 presents select frames acquired during steady-state in Cases 1 and 2. As found in the prior study by the authors, steady-state occurs after the laser has traveled approximately 1.5 mm [2]. Case 1 appears to be more consistent and shorter than Case 2. The plume is visible following above the scan track, with a slight reflection of it below the track. In contrast, the Case 2 is far less consistent.

The plume is visible following above the scan track, with a slight reflection of it below the track. In contrast, the Case 2 is far less consistent. A significant amount of spatter can be seen above and behind the track, making the images look far moreFrame 18Frame 20Frame 18Frame 16Frame 141234567891011Horizontal location (mm)1234567891011Horizontal location (mm)A) Case 1, single scan track on a bare plateB) Case 2, single scan track on a layer of powderFrame 20Frame 16Frame 14IlluminatedvaporplumeReflectionof plumeScan beginningScanendApproximatelaser spotScan direction< 550 C650 C750 C850 C950 C1038 C <Radiant temperatureSpatterFigure 2 - Select frames acquired using the thermal camera during steady state of the single scan tracks.1342Published in the Proceedings of the 2017 Annual International SFF Symposiumchaotic than the bare surface scans.Select frames during the deposition of Cases 3 and 4 are shown in Figure 3.

These frames are acquired immediately before the conclusions of their respective scan tracks. The length appears to increase as more lines are scanned and it seems that steady state is achieved shortly after the 7th scan track. In Case 4, which has a layer of powder, the amount of spatter is greater during the first scan track than in subsequent tracks.Figure 4A presents the temperature profiles extracted along the scan track from each frame acquired during steady state in Cases 1 and 2.

The vertical and horizontal lines indicate the melt pool length measurement and the solidus-to-liquidus radiant temperature [2]. For those values, one standard deviation is reported. B) an individual profile from each case to illustrate the discontinuity attributed to the solidus-to-liquidus transition.found in the earlier study are included for comparison [2]. The gray horizontal lines indicate the radiant temperature of the solidus-to-liquidus transition while the vertical gray lines indicate the front and back of the melt pool, based on the length measured in the earlier study. These plots support the observations made of Figure 2. The bare surface scan in Case 1 produces consistent temperature profiles, whereas the Case 2 scan on a layer of powder produces profiles that have a significant amount of variation and have a greater radiant temperature.

A closer investigation of these profiles suggests that, although the discontinuity associated with the solidus-to-liquidus transition cannot be detected in the powder case, it appears that the melt pool length is similar between the two cases, as illustrated with an example in Figure 4B.Figure 5 presents temperature profiles of select scan tracks first shown in Figure 3. Once again, the profiles of the bare surface scan are smooth, whereas those of the powder layer scan have a significant amount of variation, though the variation is less as the scan line number increases. Although the discontinuity associated with the solidus-to-liquidus transition is imperceptible in the powder layer radiant temperature profiles due to this variability, the melt pool length can be measured by assuming the bare surface measurements in the previous study can be applied to the powder layer.

If this assumption is valid, then in both cases the melt pool length nearly doubles in each case from the first line (approximate length of 0.76 mm) to the end of the 39th line (approximate length of 1.5 mm).The rate at which the scan track cools from 900 °C to 700 °C (radiant temperature) is calculated usingthe following equation and the results are presented in Figure 5:𝐶𝑅 =200 °𝐶 (𝑝700 °𝐶 − 𝑝900 °𝐶) 𝑟𝑣⁄where CR is the cooling rate in radiant temperature, v is the scan velocity (800 mm/s), r is the horizontal camera instantaneous field of view (0.036 mm per pixel), and p is the location (pixel) in the frame of interest where the radiant temperature profile equals 900 °C and 700 °C.

The markers for Cases 1 and 21344(2)Published in the Proceedings of the 2017 Annual International SFF SymposiumLine 1Line 3Line 7Line 39Case 3Bare plateCase 4Powder layerRadiant temperature of the solidus-liquidus transition [2]: 936 C +/-22 C55060065070075080085090095010001050-0.500.511.522.533.544.55Radiant temperature ( C)Distance from the leading edge of the melt pool (mm)Figure 5 - The radiant temperature profile near the end of select scan track that were initially presented in Figure 3. The solid horizontal line indicates the mean solidus-to-liquidus transition temperature, while the dashed horizontal lines represent one standard deviation [2].present the mean of calculations performed on all frames acquired during steady state. The standard deviation of those calculations (1σ) is reported in the text on the plot. The markers indicating Cases 3 and 4 represent the calculation from a single frame acquired at the end of each odd-numbered scan track.

The markers indicating Cases 3 and 4 represent the calculation from a single frame acquired at the end of each odd-numbered scan track. For each of the multiple scan track cases, the cooling rate decreases as the line count increases, reaching an apparent steady state value around the 20th line. The average and standard deviation (1σ) is calculated for the odd numbered scan tracks, from 21 to 39, and is reported on the figure. In both multi-line cases, the cooling rate decreases to values that are approximately 15 % to 20 % of the values measured in the single scan track cases.

The standard deviation of the measured values also decreases by an order of magnitude or more from the single scan tracks to the steady state multiple track scans.Case 2, single track, powder layer0100000200000300000400000500000600000012345678910111213141516171819202122232425262728293031323334353637383940Scan track numberCase 1, single track, bare surface0100200300400500600Cooling rate in radiant temperature ( C/s)x 103Case 4, multiple track, powder layerMultiple track average, tracks 21 through 39Case 3: 73x 103 C/s +/-3x 103 C/sCase 4: 77x 103 C/s +/-8x 103 C/sCase 3, multiple track, bare surfaceSingle scan track average during steady stateCase 1: 499 x 103 C/s +/-60x 103 C/sCase 2: 360x 103 C/s +/-100x 103 C/sFigure 6 - The rate the material cools from 900 °C to 700 °C for each odd numbered scan track.1345Published in the Proceedings of the 2017 Annual International SFF SymposiumDiscussionThe results presented in this work confirm that thermographic measurements of scan tracks on layers of powder are more complex than when the same scans are performed on bare surfaces.

Scans on powder result in greater fluctuations along the temperature profile of the scan track behind the melt pool. Considering the camera system used in this study and its resolution, these fluctuations mask the discontinuity that is associated with the solidus-to-liquidus transition and could be detected in the bare surface scans. Since the discontinuity is regularly used to identify the melt pool boundary and measure the melt pool length, the melt pool length could not be measured without assuming the radiant temperature of the transition is similar in the bare surface and powder layer scans. Further work is required to confirm if this assumption is valid. In addition, the single scan tracks on a powder layer appear to be hotter than on bare surfaces, as demonstrated in Figure 4, and the cooling rate in radiant temperature is slower, as indicated in Figure 6.

The hotter appearance and slower cooling rate in the radiant temperature scale could be attributed to two factors that require a more detailed analysis.1) The powder layer effectively increases the absorption efficiency of the laser, leading to a greater amount of energy entering the part [17] and increasing the melt pool length and the residual temperature in the track.2) The effective emissivity of the track is greater when created on a powder layer than on a bare surface, leading to an apparent hotter temperature in the radiant temperature scale even though the true temperature is equal to the bare surface scan. The effective emissivity could be elevated in the powder layer cases by: a. The rougher surface of the scan track [18]. b. The elevated scan track (due to the solidified powder layer) changing the relativeviewing angle, which has been shown to affect emissivity [12], [13].c.

b. The elevated scan track (due to the solidified powder layer) changing the relativeviewing angle, which has been shown to affect emissivity [12], [13].c. Reflections from the spatter on the scan track, increasing the signal measured by thecamera according to Equation (1).A more detailed investigation is required to understand the impact of each of these possible factors to understand whether powder does impact the melt pool length and the temperature in single track scans, or if it only appears to have an effect.Analysis of the multiple scan track cases suggest that the temperature between the bare surface and powder layer scans are not as different as they appear in the single line cases. As the scan lines count increases, the temperature profiles of the two cases have better agreement, as shown in Figure 5. Furthermore, the variability appears to decrease, as does the spatter, after the first few lines, as indicated in Figure 5 and Figure 6.

These observations suggest the following:1) Spatter can have a significant impact on both the magnitude of the temperature profile and its variability.2) The deposition of successive tracks, that are deposited moving away from the camera, has a different impact on effective emissivity than that of a single scan track.3) The difference in energy absorption between scans on bare surface and single layers of powder is only significant for the first layer, since denudation draws powder into the track, exposing the bare surface for each successive track [19], [20].Once again, a more detailed experiment must be designed to investigate these and other possibilities.1346Published in the Proceedings of the 2017 Annual International SFF SymposiumConclusionsThis work presented thermographic measurements acquired during the powder bed fusion process. Single and multiple scan tracks were executed on both bare surfaces and hand-spread powder layers.

Single and multiple scan tracks were executed on both bare surfaces and hand-spread powder layers. The results were compared to gain insight into the effect of powder on both the process and the measurement. Results indicate that the presence of powder increases the variability in the radiant temperature profiles behind the melt pool that that while the single scan tracks appear to be hotter in radiant temperature, it is unclear if the increase is due to differences in true temperature or the effective emissivity of the surface. The multiple scan track results indicate that the bare surface and powder layer scans are more similar as the number of scan lines increase. Future work will improve upon the experimental method and analysis to answer questions that could not be answered in this preliminary study.References[1] M. Doubenskaia, A. Domashenkov, and I.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of Powder Particle Size on the Fabrication of Ti-6Al-4V using Direct Laser Metal Deposition from Elemental Powder Mixture Xueyang Chen1, Lei Yan1, Wei Li1, Frank Liou1, and Joe Newkirk21Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409 2Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractDirect Laser Metal Deposition (LMD) was used to fabricate thin-wall Ti-6Al-4V using the powder mixture of Ti-6 wt.%Al-4 wt.%V.

Scanning electron microscopy (SEM), optical microscopy (OM) and energy dispersive spectroscopy (EDS) were employed to examine the chemical composition and microstructure of the as-deposited sections. Vickers hardness tests were then applied to characterize the mechanical properties of the deposit samp les which were fabricated using pre-mixed elemental powders. The EDS line scans indicated that the chemical composition of the samples was homogenous across the deposit. X-ray diffraction (XRD) was used for the phase identification. After significant analysis, some differences were observed among two sets of deposit samples which varied in the particle size of the mixing Ti-6wt.%Al-4wt.%V powder. It could be found that the set with similar particle number for Ti, Al and V powder made composition much more stable and could easily get industry qualified Ti-6Al-4V components.1. IntroductionLaser Metal deposition is an additive manufacturing technique.

IntroductionLaser Metal deposition is an additive manufacturing technique. The applications of this technique include coatings, rapid prototyping, tooling, and refurbishment. As shown in Figure 1, laser deposition uses a focused laser beam as a heat source to create a melt pool on an underlying substrate. Powder material is then injected into the melt pool through nozzles. The incoming powder is metallurgically bonded to the substrate upon solidification [1-7]. Conventionally, Direct Laser Deposition (DLD) manufactured parts by adding materials layer by layer using pre-alloyed powder. In the pre-alloyed powder, each individual powder particle has the composition o f the desired alloy composition in the final part [8]. But sometimes, it is more flexible to use the mixed elemental powder to develop products.

But sometimes, it is more flexible to use the mixed elemental powder to develop products. In the elemental powder mixture, each individual powder particle only has the composition of a single element present in the desired final alloy composition, but the blend of the total elemental powder forms the desired final alloy composition [8]. In this study, the material of both powder and substrates is Ti-6Al-4V alloy, which is widely used in the aerospace industry and especially suitable for laser processing. Owing to high cost, this titanium alloy is usually developed by near-net-shape Powder Metallurgy (PM) methods such as laser deposition. Ti-6Al-4V is still classified as one of the extremely difficult-to-machine materials using conventional machining which makes DLD an ideal manufacturing process for the same [9].714Figure 1. Schematic of the MST-LAMP DLD system [10]2. Experimental proceduresElemental powders used in this study were titanium, aluminum, and vanadium.

Schematic of the MST-LAMP DLD system [10]2. Experimental proceduresElemental powders used in this study were titanium, aluminum, and vanadium. The elemental powders were characterized to confirm information provided by suppliers and also to determine particle shape and size. By understanding the size distribution and shape of particles, observations and any findings in deposits could potentially be correlated with particle shape or size. Images were taken using a Hitachi S4700 SEM and image analysis was performed using the ImageJ software.Figure 2 shows the SEM images and size distribution of Titanium, Aluminum and Vanadium powder. Titanium powder, grade Ti-109, was purchased from Atlantic Equipment Engineers for use in deposition of Ti-6Al-4V alloys. This powder, grade Ti-109, was listed as being 99.7% and of -100 mesh (<149 μm) size. A SEM micrograph of the titanium powder can be seen in Figure 2(a) and indicates significant complexity in particle shape.

A SEM micrograph of the titanium powder can be seen in Figure 2(a) and indicates significant complexity in particle shape. This makes it difficult for the automatic particle analysis software in ImageJ to accurately outline individual particles and calculate their surface area. While this analysis may not give precise measurements, trends in particle size and distribution can be observed. From the resulting software analysis, an average particle size of 106 μm was determined. The elemental aluminum powder, grade Al-103, was produced by Atlantic Equipment Engineers. The purity of the powder was listed as 99.8%. Using similar software analysis gave the Al-103 an average particle size of 165 μm. The elemental vanadium powder, grade VA-103, was purchased from the Atlantic Equipment Engineers with a powder purity of 99.8%.

The elemental vanadium powder, grade VA-103, was purchased from the Atlantic Equipment Engineers with a powder purity of 99.8%. A SEM micrograph of the vanadium powder can be seen in Figure 2(c) and the average size of this powder was 175 μm.715a) T itanium powder SEM imageb) Aluminum powder SEM imagec) Vanadium powder SEM imaged) Particle size distributionFigure 2. The SEM images and size distribution of Titanium, Aluminum and Vanadium powder [10]Before the deposition process began, the elemental powder mixes were prepared. Powders were weighed out according to desired alloy composition in a glove box under an argon atmosphere and placed in a sealed bottle. The powder mixture was a representative of the Ti-6Al- 4V alloy. After the desired composition was obtained, the bottle was placed in a Turbula® mixer for 20 minutes. A 1KW Nd:Yag laser was used in this study to deposit these elemental alloy powders.

After the desired composition was obtained, the bottle was placed in a Turbula® mixer for 20 minutes. A 1KW Nd:Yag laser was used in this study to deposit these elemental alloy powders. The laser has a wavelength of 1024 nm and approximate 1mm spot size at a 750 mm focal length. To ensure uniformity, thin wall samples with 30 mm track length were deposited on Ti-6Al-4V substrates during the study. Also, all the substrates were 2 inches long, 0.5 inch wide, and 0.25 inch thick. Before deposition, a preheating process was performed by using the laser across the surface of the substrate with 1kW power and without powder. The preheating process was used to form a melt pool on the substrate and clean the surface of the substrate off of scale buildup and other impurities and surface imperfections.

The preheating process was used to form a melt pool on the substrate and clean the surface of the substrate off of scale buildup and other impurities and surface imperfections. This research theorizes that this pre-heat and clean-up will enable conductive metallurgical bonding of the powders with the substrates.The commercially irregular pure Ti, Al, and V powders were weighed out and two sets of powder blends were prepared to facilitate similar particle acceleration and similar particle number. It is believed that this will enable a relatively uniform shape of the deposit from top to bottom716using the laser deposition system. Ti, Al, and V powders were mixed with the designated weight percentage ratio of 90:6:4 to get target material Ti-6Al-4V. Laser transverse speed was set at 1000 mm/min for all the paths.

Ti, Al, and V powders were mixed with the designated weight percentage ratio of 90:6:4 to get target material Ti-6Al-4V. Laser transverse speed was set at 1000 mm/min for all the paths. For laser power, the high normal value of 1 kW was first selected to generate melt pool and then laser power was dropped to and kept stable at 450 W (after 10 layers) to maintain a stable melt pool. While depositing the material, uniformity was ensured in the deposition parameters of travel speed, powder feed rate, and layer thickness constant from test to test. This would lead to an unbiased comparison between different depositions by using similar powder particle and similar powder particle number. Using the analysis, the research theorizes that optimized parameters can be determined in the results for different deposits.During the laser deposition process, powder blends with the designated weight percentage ratio were injected into the melt pool through the nozzle.

To confirm to get target material with the original weight percentage ratio, each element in the blend with the desired composition needs to be melted at the same time. The powder particles had the same initial speed when they went into the nozzle. So particle acceleration is a key factor for composition control. According to the fluid mechanics [11], the particle acceleration is inversely proportional to the particle density and the square of powder diameter, which could be expressed as 𝑎 ∝ 1/(𝜌 ∗ 𝐷 2). Also, particle size for each element in powder blend has a relationship with the particle number. So particle size of each element could be changed to control the composition in the final part. Then the hypothesis proposed in this research is proved using experiments design around two sets of data namelySet 1: A comparative study of the acceleration of the elemental powder particles (Table 1).

Set 2: A comparative study of the number of particles in each elemental powder (Table 2).Among the two sets of experiments conducted, the appropriate particle size of Ti, Al and V powder has been selected according to the particle diameter listed in Table 1, it shows the particle size and calculated acceleration ratio of Ti, Al and V powder. In this experiment, Ti, Al, V elemental powder has the similar particle acceleration during the process. Similarly, after choosing suitable particle size for set 2, corresponding particle number ratio of Ti, Al, V powder could be calculated in Table 2. In this set, Ti, Al, V elemental powder has the similar particle number during the process. These two sets of experiments were chosen to study which factor, among particle acceleration and particle number, would affect an as-built part’s final chemistry composition.Table 1.

These two sets of experiments were chosen to study which factor, among particle acceleration and particle number, would affect an as-built part’s final chemistry composition.Table 1. Particle Size and Acceleration Ratio in Set 1Irregular powder Ti Al VDiameter 112 133 96Acceleration ratio 1 1.18 1.06Table 2. Particle Size and Number Ratio in Set 2Irregular powder Ti Al VDiameter 147 94 63Number ratio 1 0.43 0.44717After deposition, longitudinal samples were cut off using the EDM machine and mounted for grinding and polishing. Kroll’s Reagent consisting of 92 ml distilled H2O, 6 ml HNO3, and 2 ml HF was used for etching. To research the microstructure characteristic and composition in the elemental powder deposits, energy dispersive spectroscopy (EDS) line scans were performed using Hitachi S4700 and FEI Helio 600 scanning electron microscope (SEM). Also, optical microscopy was used to assess the microstructure in the samples.

Also, optical microscopy was used to assess the microstructure in the samples. Comparison of the mechanical properties between the two sets of different samples was also conducted. The Vickers hardness measurements were taken with a Struers Duramin micro-hardness tester. 9.8kg load with a dwell time of 10 seconds was applied during the test hardness procedures.3. Results and discussion3.1 Vickers hardness & EDS testFigure 3 shows the composition distribution and Vickers hardness of the deposited sample from elemental powders of Set 1 whose three elements of Ti, Al and V have the similar particle acceleration. Ti-6Al-4V substrate Vickers hardness was first taken near the bottom edge that is away from heat affected zone (HAZ) and found to be 352.44±15.41 HV. It was observed that the Vickers hardness values increased as it progressed into the HAZ from the substrate.

It was observed that the Vickers hardness values increased as it progressed into the HAZ from the substrate. In the deposit zone, there was a temporary increase for the first several layers, then a decrease and stabilization from the middle to the top of the deposit.For the deposited samples using mixed powder of Set 1, area scan and line scan were employed to investigate the composition distribution along the build height direction. Figure 3 shows the EDS result of the deposit. The weight percentage of Ti, Al and V could be found in the image in the left figure of Figure 3. Al and V curves were focused on and showed their weight percentage variation against the position in the image on the right. Two marked zones in this figure mean industry qualified Al and V weight percentage for Ti-6Al-4V, which is Al ranges from 5.5- 6.75% and V ranges from 3.5-4.5%.

Two marked zones in this figure mean industry qualified Al and V weight percentage for Ti-6Al-4V, which is Al ranges from 5.5- 6.75% and V ranges from 3.5-4.5%. At deposit height, Al and V weight percentage showed that the values varied outside the qualified range, which means the as-deposit alloy was not the targeted material Ti-6Al-4V. There was some powder loss when the powder mixture went through the nozzle, so it was supposed that less powder in the blend especially for Al and V powder because of the low volume of them. This would lead to the result that the weight percentage of Al and V along build height direction vary a lot.718Figure 3. Composition distribution & Vickers Hardness along build height direction of Set 1Figure 4 shows the composition distribution and Vickers hardness of the deposited sample from elemental powders of Set 2 whose three elements of Ti, Al and V have similar particle number.

Also, Ti-6Al-4V substrate Vickers hardness was first taken near the bottom edge that is away from heat affected zone (HAZ) and found to be 359.11±13.78 HV. In this case, it was observed that the Vickers hardness values decreased with progress into the HAZ from the substrate side. In the deposit zone, there was an increase for the first several layers and stabilization beyond that.For the newly deposited samples using mixed powder (Set 2), area scans and line scans were also employed to check the composition distribution along the build height direction. Figure 4 shows the EDS results of the deposits. The weight percentage of Ti, Al and V could be found in the image on the left of Figure 4. Al and V curves were focused on and showed their weight percentage variation in the image on the right of Figure 4. Two marked zones in the right figure mean industry qualified Al and V weight percentage for Ti-6Al-4V.

Two marked zones in the right figure mean industry qualified Al and V weight percentage for Ti-6Al-4V. It can be seen that the deposit had qualified Al and V weight percentage and get close to the maximum limit of industry qualified Ti-6Al-4V, which means industry qualified Ti-6Al-4V has been made. From this, it can be concluded that powder number for each of the elements in the blends is a key factor towards the composition acquired in the final part. It could be found that similar particle number made composition much more stable and close to industry qualified Ti64. The research theorizes that the increase in the volume of Al and V in the set 2’s powder blends increases the probability of those two elements particles falling into melt pool.719Figure 4.

The research theorizes that the increase in the volume of Al and V in the set 2’s powder blends increases the probability of those two elements particles falling into melt pool.719Figure 4. Composition distribution & Vickers Hardness along build height direction of Set 2Figure 5(a) showed the line scan of the deposited sample from elemental powders of Set 1 whose three elements of Ti, Al and V have similar particle acceleration. Figure 5(b) showed the line scan of the deposited sample from elemental powders of Set 2 whose three elements of Ti, Al and V have similar particle number. Line scan was shown in Figure 5(a) to define composition homogeneity along the build height direction. Thus, it can be known that similar particle acceleration could not guarantee a homogeneous chemistry distribution throughout the deposit from Figure 5(a).

Thus, it can be known that similar particle acceleration could not guarantee a homogeneous chemistry distribution throughout the deposit from Figure 5(a). Also, Line scan was shown in Figure 5(b) to define composition homogeneity along the build height direction which shows that the distribution is homogenous for the deposit.a) Similar particle accelerationb) Similar particle numberFigure 5. Line scan of the deposited samples7203.2 MicrostructureFigure 6 shows the microstructure of deposited Ti-6Al-4V from elemental powder mixture with the similar particle acceleration at the same laser power and laser transverse speed. The optical microscope images captured from the bottom part, middle part and top part of the melt pool (Figure 6 (a) to (c)) in the deposit contained basketwave, colony and martensite microstructure respectively. Figure 6(a) showed the finer microstructure could be observed in the bottom part compared with the middle and top part of the deposit.

Figure 6(a) showed the finer microstructure could be observed in the bottom part compared with the middle and top part of the deposit. Also, a steady state could be observed from middle part to top part and there were limited differences in the microstructure between them. In these two zones, the Widmanstätten Basketwave is the main type of microstructure that could be found. It tends to form with an increasing cooling rate from the β-phase field [12-13].Analyses of these microstructures revealed that the smaller grain sizes were achieved closer to the substrate in the deposit. The reason was the proximity to the substrate, increased the cooling rate, which would lead to a finer microstructure [14]. The same phenomenon was observed in the microstructure of the deposited samples from the elemental powder mixture with the similar particle number of Set 2.

The same phenomenon was observed in the microstructure of the deposited samples from the elemental powder mixture with the similar particle number of Set 2. Therefore, it was concluded that the steady state section tends to have the lower cooling rate, which would lead to the formation of coarser microstructure [15]. Figure 7 shows the microstructure of deposited Ti-6Al-4V from elemental powder mixture with the similar particle number using the same laser power and laser transverse speed. Also, a steady state could be observed from middle part to top part and Widmanstätten Basketwave is the main type of microstructure that could be found.a) Bottom of the depositb) Middle of the depositc) Top of the depositFigure 6. Microstructure of the deposits using elemental powders of Set 1a) Bottom of the depositb) Middle of the depositc) Top of the depositFigure 7. Microstructure of the deposits using elemental powders of Set 27214.

Microstructure of the deposits using elemental powders of Set 27214. ConclusionsThe blended elemental powder can be used in the direct laser deposition process to offer a variety of cost-effective possibilities for near-net-shape manufacture and functionally graded materials. The finer microstructure could be observed in the bottom part of the deposit compared with the steady state in the middle and top part of the deposit. The Vickers hardness could fluctuate at the first several layers of the deposit and then stayed stable at the top layers of the deposit. Particle number for each element in powder blends is a key factor for composition control in the final part. Similar particle number for Ti, Al and V powder could easily get industry qualified Ti- 6Al-4V components.5. References[1] Chua, C.K., Leong, K.F., Lim, C.S., Rapid Prototyping: Principles and Applications, 3rd Edition, 2010. [2] Unocic, R.R.

Powder reuse may adversely affect the powder characteristics, including the flowability, size distribution, chemical composition, resultant microstructural, and consequently, mechanical properties of the fabricated parts. This study aims to investigate the effect of powder reuse on the microstructural features and fatigue performance of Ti-6Al-4V specimens fabricated using a directed energy deposition (DED) process. Characteristics of reused powder particles, such as the size distribution and chemical composition, were evaluated and compared with that of virgin powder. Microstructural features and characteristics of the process-induced defects were examined using scanning electron microscopy and x-ray computed tomography, respectively.

Microstructural features and characteristics of the process-induced defects were examined using scanning electron microscopy and x-ray computed tomography, respectively. Fatigue performance of the specimens fabricated using reused powder was evaluated and compared to their control counterparts, fabricated using virgin powder.Keywords: Additive manufacturing; Powder recycling; Virgin powder; Powder flowability; Powder size distribution5801. IntroductionThe domain of metal additive manufacturing (AM) contains different techniques that join materials in a layer-wise manner, providing an opportunity to produce complex-shaped objects directly from 3D model data [1,2]. Laser-direct energy deposition (LDED), also known as laser- engineered net shaping (LENS®), is a powder-based metal AM process commonly used for fabricating metallic parts [3]. In this technique, metal powder is blown into the melt pool and consecutive layers are fabricated along the build direction [4].

In this technique, metal powder is blown into the melt pool and consecutive layers are fabricated along the build direction [4]. Theoretically, blown powder particles should be delivered to the melt pool zone, however, some of the dispensed powder particles, experiencing laser heating, may not get captured in the melt pool and hit the cold surface of the build [5]. Low deposition efficiency of the LDED process results in large amounts of unused powders, which is economically and environmentally expensive [1,6]. Therefore, the leftover powder is usually reused/recycled for future builds.Experiencing laser heating, the leftover powder particles may experience changes in their characteristics, such as powder flowability, size distribution, chemical composition, etc. [7,8].

[7,8]. Changes in the characteristics of powder may affect the microstructural features and formation of process-induced defects, i.e., gas pores and lack-of-fusion (LOF) defects, leading to variations in the mechanical performance of the fabricated parts, particularly their fatigue performance. Furthermore, due to the high reactivity of some materials like titanium alloys with oxygen at high temperatures, oxygen contamination during the powder recycling process can be detrimental to the quality and structural integrity of the part [9].

Therefore, there is a need to understand and correlate the effects of powder reuse/recycling on the microstructural and mechanical performance of the fabricated part.Studies on Inconel 718 [5,10], 316L stainless steel (SS) [11], and Ti-6Al-4V [9] alloys fabricated via LDED process showed that powder recycling causes little variations in flowability [5,11], discoloration in microstructures [11] or powder particles [10], but no significant changes in chemical compositions of powder [5,9–11], oxygen content [5,9], and microhardness values [5,10]. While some studies reported an increase in particle size distribution due to powder reuse [9–11], an opposite trend was reported for particle size distribution by Carroll et al. [5]. Carroll et al. reported a marginal decrease in particle size as a result of reusing the Inconel 718 powder [5]. Terrassa et al.

[5]. Carroll et al. reported a marginal decrease in particle size as a result of reusing the Inconel 718 powder [5]. Terrassa et al. reported that recycling of powder had no significant effect on compression yield stress and tensile properties of 316L stainless steel (SS) specimens [11].A review of the published literature shows that powder recycling is a trending topic in powder bed fusion (PBF) AM systems. Several research efforts have been done aiming to understand how powder recycling influences the characteristics of PBF parts [12–20]. However, there are only a few research efforts focusing on how the reused powder feedstock influences the properties of parts fabricated via LDED process [5,9–11,21]. As a result, there is very limited knowledge on how reused powder may affect the structural integrity of LDED parts.

As a result, there is very limited knowledge on how reused powder may affect the structural integrity of LDED parts. Although results of current research on powder recycling in PBF processes may provide useful insights, observations and outcomes of these studies cannot be directly applied to the parts fabricated via DED process, considering the fact that the characteristics of the build process in PBF systems being different from that of DED systems. To fill this gap, the present work aims to investigate the effect of powder reuse on the characteristics of the recycled powder (i.e., oxygen content, chemical581composition, particle size distribution, and powder flowability) as well as the microstructure and fatigue performance of LDED Ti-6Al-4V.2. Experimental Procedure2.1.

Experimental Procedure2.1. Specimen fabrication and preparationPlasma atomized Ti-6Al-4V powder with a particle size distribution of 45-90 μm (LPW technology) was utilized in this study to fabricate samples via an OPTOMEC LENS 750 system equipped with a 1 kW laser source (Nd:YAG). As-received (i.e., virgin) and recycled (i.e., un- melted powder particles recovered by previous depositions) powders were used as starting feedstock material to fabricate two groups of fatigue specimens: (i) virgin and (ii) recycled. The recycled powder was obtained through nine deposition cycles. To remove any under/over-sized particles after each deposition, the powder collected from the build stage was sieved, and the agglomerates larger than 150 µm, and particles smaller than 45 μm were removed. Before starting the deposition in each cycle, the build plate was prepared by sandblasting and cleaning with soap and alcohol.

Before starting the deposition in each cycle, the build plate was prepared by sandblasting and cleaning with soap and alcohol. Figure 1 shows the flowchart of the procedure utilized for the collection of recycled powder.Figure 1. Flowchart of the building procedure and collecting recycled powder.Process parameters utilized for the fabrication of Ti-6Al-4V samples, based on previous operating experience with the alloy [22,23], are given in Table 1. Vertically oriented cylindrical coupons with 85.0 mm in height and 8.0 mm in diameter were fabricated using virgin and recycled powders. Arrangement of the cylindrical coupons on the build plate is shown in Figure 2(a). All samples were annealed at 760 °C for 1 hour in an inert Argon atmosphere, followed by air cooling. This procedure allows for stress relief and decomposition of the martensitic phase [24].

All samples were annealed at 760 °C for 1 hour in an inert Argon atmosphere, followed by air cooling. This procedure allows for stress relief and decomposition of the martensitic phase [24]. Cylindrical rods were machined to the dogbone shape, with a reduced diameter of 4.0 mm at the gage section, as shown in Figure 2(b), for uniaxial fatigue testing in accordance with ASTM E466 [25].Table 1. Process parameters utilized in this study for fabrication of the Ti-6Al-4V samples.Process ParametersLaser power (W) Travel speed (mm/min) Powder flow rate (rpm) Hatch distance (μm) Hatch rotation (degree) Scan pattern Layer thickness (μm)350 1016 2.5 508 120 Zig-zag 508582Figure 2. Arrangement of the cylindrical coupons on the build plate.2.2. Material characterizationIn order to evaluate the powder characteristics, including the chemical composition, particle size distribution, and flowability, a 50 g sample of powder was collected from the virgin powder and recycled powder.

Powder morphology, microstructural features, and process-induced defects were examined using scanning electron microscopy (SEM). Energy dispersive X-ray analysis (EDAX) was used to identify the elemental composition of powder samples. The oxygen content of the powder samples was analyzed using a LECO ONH836 elemental analyzer. Particle size distribution of powder samples was measured by a laser analyzer (Horiba LA 950). The flowability of powders was measured using a Hall flowmeter according to the ASTM B213 [26]. In order to evaluate the formation of any possible process-induced defects, fatigue specimens were scanned using an X-ray computed tomography (XCT) system. To improve scanning resolution, only the gage section of specimens and each specimen separately was scanned. Uniaxial fully- reversed (R = -1) fatigue tests were performed under the load-controlled mode using a servo- hydraulic load frame at room temperature.

Uniaxial fully- reversed (R = -1) fatigue tests were performed under the load-controlled mode using a servo- hydraulic load frame at room temperature. Fatigue tests were conducted using a sinusoidal loading waveform until failure occurred or 107 cycles were achieved, in which the test was considered to be a ‘run-out’.3. Results and Discussion3.1. Powder characterizationResults of oxygen content analysis revealed a slight increase in the concentration of the oxygen for the recycled powder as compared to their virgin counterparts. As shown in Figure 3, powder recycling caused 18% increase in the amount of oxygen, as the oxygen content from 0.076583wt.% in the virgin powder raised to 0.089 wt.% in the recycled powder. Although the increase in the amount of oxygen was noticeable, the overall oxygen content did not exceed the allowable oxygen content (i.e., 0.2 wt.%) for the Ti-6Al-4V, specified by the ASTM F2924-14 standard [27].

The EDAX analysis did not show significant differences in the elemental composition of titanium between virgin and recycled powders, as shown in Figure 4. However, due to powder recycling, a noticeable loss in aluminum (14%) and a rise in vanadium (21%) were observed. Our observation regarding the increase in vanadium concentration is in contrary to what was reported by Rousseau et al. [9], in which a loss of vanadium in the reuse powders was observed.Figure 3. Oxygen content in the virgin and recycled powders.Figure 4. Elemental composition of the virgin and recycled powders (atomic percentage of titanium, aluminum, and vanadium) obtained by EDAX analysis.In order to observe the changes in the powder particle morphology after recycling, the virgin and recycled powder samples were examined using SEM. The SEM images of particle584morphology of the virgin powder and recycled powder are shown in Figures 5(a) and (b), respectively.

The SEM images of particle584morphology of the virgin powder and recycled powder are shown in Figures 5(a) and (b), respectively. Observations indicated that the recycling process caused deformation in the shape of the particles and increase in the number of satellite particles sticking to the surface of larger particles. The particle shape and the presence of satellite powder will seriously affect the flowability and other properties of the powder. The presence of satellite powder particles was noticeably increased after powder recycling, as seen by comparing Figure 5(a) with (b). The formation of satellite particles was most likely due to particle collisions during the deposition process. In addition, some powder particles were partially melted and stacked to other powder particles during the deposition process, forming larger agglomerated particles, as can be seen in Figure 5(b).Figure 5.

SEM images of the (a) virgin powder and (b) recycled powder.The size distribution for the virgin and recycled powder is presented in Figure 6. As can be seen, size distribution of the recycled powder is skewed as compared to the virgin powder. In addition, the mean size of particles after recycling has noticeably increased, agreeing with the results reported by Renderos et al. [10]. Particle size distribution can influence the powder flow rate, which must be controlled for the quality of the fabricated parts [11]. Despite the noticeable differences observed in the particle size distribution and morphology of virgin and recycled powders, results from the flowability analysis revealed a slight difference in the flowability of recycled powders compared with virgin counterparts. Table 2 presents the flow rates of the virgin and recycled powders measured by the Hall flowmeter according to ASTM B213 [26].

Table 2 presents the flow rates of the virgin and recycled powders measured by the Hall flowmeter according to ASTM B213 [26]. As can be seen, the flow time of recycled powder is slightly higher as compared to virgin powder.585Figure 6. Particle size distribution of the virgin and recycled powders.Table 2. Flow rates of the virgin and recycled powders, measured by the Hall flowmeter according to ASTM B213 [26].Virgin PowderRecycled PowderElapsed time (sec)20.2321.883.2. Microstructural characterizationThe SEM micrographs of the samples fabricated using virgin and recycled powders are presented in Figure 7. No significant differences in the grain size and morphology were observed, most likely due to the fact that both groups underwent the annealing treatment at 760 °C for 1 hour. A discoloration was observed on recycled samples, similar to the observation reported by Renderos et al. and Terrassa et al. [10,11].Figure 7.

A discoloration was observed on recycled samples, similar to the observation reported by Renderos et al. and Terrassa et al. [10,11].Figure 7. SEM micrographs of the samples fabricated using (a) virgin powder and (b) recycled powder.586Results from the XCT analysis, including the number of defects, largest volume of defects, average volume of defects and total volume of defects, are presented in Figure 8. Each value was obtained by getting average of all specimens in their category. For example, the largest defect volume for the virgin category represents the average of largest defect volume detected by XCT in the specimens fabricated via virgin powder. As can be seen in Figure 8(a), the number of defects in specimens fabricated using recycled powder is significantly higher as compared to their virgin counterparts, i.e., fabricated using virgin powder.

This can be attributed to the formation of process-induced defects, i.e., gas pores and lack of fusion (LOF) defects, resulting from changes in characteristics of powder during recycling. As can be seen in Figures 8(b) and (c), the largest and average volume of detected defects in the specimens fabricated using recycled powder is almost doubled as compared to their virgin counterparts. Comparing the increase observed in the number of defects with the average volume of defects in the recycled specimens indicates that the huge rise in the total volume of defects should be more attributed to the increase in the number of defects. Although the microstructure observations showed no notable differences, the XCT results revealed a significant difference in the level of process-induced defects between the specimens fabricated using virgin powder and recycled powder.Figure 8.

XCT analysis of defects for the specimens fabricated using virgin and recycled powders: (a) number of detected defects, (b) largest defect volume, (c) average defect volume, and (d) total volume of detected defects.5873.3. Fatigue behaviorFigure 9 presents the uniaxial fully-reversed (R = -1) fatigue test results for the LDED Ti- 6Al-4V specimens fabricated using virgin and recycled powders. The specimens fabricated using virgin powder showed a higher fatigue resistance than those fabricated using recycled powder in both short-life and long-life regimes. The lower fatigue resistance of the recycled specimens as compared to their virgin counterparts is directly attributed to the higher level of defects in recycled specimens.Fractography of fatigue fracture surfaces revealed that the process-induced LOF defects were the main source of crack initiation in both groups of fatigue specimens, i.e., fabricated using virgin powder and recycled powder, as shown in Figure 10.

In general, defects that initiated cracks in the recycled specimens appeared to be larger, resulting in shorter fatigue lives of the recycled specimens as compared to their virgin counterparts. In most cases, LOF defects that served as the crack initiation sites were attached to the surface, as can be seen in Figure 10(b). However, in some cases, subsurface LOF defects were observed to be the crack initiation site. Due to cyclic loading, these defects extend to the surface, indicating that subsurface LOF defects still play a dominant role in the fatigue life of the specimens.Figure 9. Stress-life fatigue data of the LDED Ti-6Al-4V specimens, fabricated using virgin and recycled powders.588Figure 10. (a) An overall view of a fracture surface for fatigue specimens fabricated using recycled powder at high stress amplitude, i.e., 300 MPa, (Nf = 23,441 cycles), and (b) a magnified view of the LOF defect served as the crack initiation site.4.

SummaryIn this study, the effects of powder reuse/recycling on the powder characteristics as well as the microstructure and fatigue behavior of LDED Ti-6Al-4V samples were investigated, and the results were compared with their counterparts fabricated using virgin powder. A slight increase was observed in the contents of oxygen, titanium, and vanadium for the powder after nine times of recycling. The recycled powder had a noticeably smaller particle mean size as compared to the virgin powder. The flowability of the recycled powder, measured using the Hall flowmeter, was slightly lower than that of the virgin powder. The XCT results revealed a significant increase in the number and size of process-induced defects for the specimens fabricated using recycled powders. Containing a higher level of porosity, the specimens fabricated using recycled powder showed a lower fatigue resistance compared to their counterparts fabricated using virgin powder.

Containing a higher level of porosity, the specimens fabricated using recycled powder showed a lower fatigue resistance compared to their counterparts fabricated using virgin powder. For both groups of fatigue specimens, i.e., fabricated using virgin powder and recycled powder, process-induced LOF defects served as the crack initiation site.5. References[1]Thompson SM, Bian L, Shamsaei N, Yadollahi A. An overview of Direct Laser Deposition for additive manufacturing; Part I: Transport phenomena, modeling and diagnostics. Addit Manuf 2015;8:36–62.[2] ASTM Standard F2792-12a: Standard Terminology for Additive ManufacturingTechnologies 2015.[3] Qiu C, Ravi GA, Dance C, Ranson A, Dilworth S, Attallah MM. Fabrication of large Ti–6Al–4V structures by direct laser deposition. J Alloys Compd 2015;629:351–61.[4] Yadollahi A, Mahtabi MJ, Khalili A, Doude HR, Newman JC.

Effect of the reuse of powder on material properties of Ti6Al4V processed by SLM. AIP Conf Proc 2019;2113:150006.[20] Nezhadfar PD, Soltani-Tehrani A, Sterling A, Tsolas N, Shamsaei N. The Effects of Powder Recycling on the Mechanical Properties of Additively Manufactured 17-4 PH Stainless Steel 2018.[21] Rousseau JN, Bois-Brochu A, Blais C. Effect of oxygen content in new and reused powder on microstructural and mechanical properties of Ti6Al4V parts produced by directed energy deposition. Addit Manuf 2018;23:197–205.[22] Marshall G, Young WJJ, Shamsaei N, Craig J, Wakeman T, Thompson SM. Dual Thermographic Monitoring of Ti-6Al-4V Cylinders During Direct Laser Deposition 2015.590[23] Marshall GJ, Thompson SM, Shamsaei N. Data indicating temperature response of Ti–6Al– 4V thin-walled structure during its additive manufacture via Laser Engineered Net Shaping. Data Brief 2016;7:697–703.[24] Zhai Y, Galarraga H, Lados DA.

Data Brief 2016;7:697–703.[24] Zhai Y, Galarraga H, Lados DA. Microstructure, static properties, and fatigue crack growth mechanisms in Ti-6Al-4V fabricated by additive manufacturing: LENS and EBM. Eng Fail Anal 2016;69:3–14.[25] ASTM Standard. E466—standard practice for conducting force controlled constantamplitude axial fatigue tests of. Metallic Materials. 2015.[26] ASTM B213-20. Standard Test Methods for Flow Rate of Metal Powders Using the HallFlowmeter Funnel 2020.[27] ASTM F2924 - 14 Standard Specification for Additive Manufacturing Titanium-6Aluminum-4 Vanadium with Powder Bed Fusion. 2021.591.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of Preheating Build Platform on Microstructure and Mechanical Properties of Additively Manufactured 316L Stainless Steel P.D. Nezhadfar1,2, Arash Soltani-Tehrani1,2, Nima Shamsaei1,21 Department of Mechanical Engineering, Auburn University, Auburn, AL 36849, USA 2 National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849, USAAbstractThis study aims to understand the effect of build platform preheating on the microstructural features and mechanical properties of 316L stainless steel (SS) fabricated via laser beam powder bed fusion (LB-PBF) process. Two sets of specimens were fabricated on a non-preheated build platform and the build platform preheated to 150 °C.

Two sets of specimens were fabricated on a non-preheated build platform and the build platform preheated to 150 °C. Thermal simulations are carried out using ANSYS using additive manufacturing module to investigate the variation in thermal history experienced by the specimens in each condition. Microstructural features are analyzed via simulation, and the results are validated experimentally. In addition, the effect of preheating on the porosity size and distribution is evaluated using digital optical microscopy.

In addition, the effect of preheating on the porosity size and distribution is evaluated using digital optical microscopy. Mechanical properties of specimens from each condition are further assessed and correlated to the variations in microstructure and defect size distributions.Keywords: Laser beam powder bed fusion (LB-PBF); Preheating; Stainless steel; Porosity; Mechanical properties; MicrostructureIntroductionEstablishing process-structure-property-performance (PSPP) relationships for various material systems fabricated via additive manufacturing (AM) techniques has been becoming of particular interest for various industrial sectors. Although AM provides an opportunity to manufacture parts with complex geometries, the intricate thermal history experienced by the parts during fabrication leads to the formation of defects (i.e. pores, lack of fusion, surface roughness), producing anisotropic microstructure [1,2] as well as inducing residual stress.

pores, lack of fusion, surface roughness), producing anisotropic microstructure [1,2] as well as inducing residual stress. Daniewicz and Shamsaei [1] introduced the fatigue and fracture of additively manufactured parts as the most challenging issue hindering the adoption of AM for critical load bearing components. Several other studies also have shown that the presence of the aforementioned defects reduces the fatigue resistance of AM parts. Accordingly, as the most critical load-bearing parts are usually experiencing cyclic loading in service, it is essential to reduce defects and microstructure heterogeneity by controlling the thermal history.Thermal history is controlled by the employed process parameters (i.e. laser power, scan speed, scan strategy, etc.) [3], part size and geometry [4], build orientation [5], and the fabrication environment (due to shielding gas thermal properties and enclosure walls) [6].

These have been reported as the controlling parameters for reducing the defects and anisotropy in the microstructure [2] and residual stress. Moreover, controlling thermal history may eliminate the need for post- processing procedures (e.g. heat treatment, surface machining, hot isostatic pressing (HIP)), which expedites the adoption of AM for widespread industrial applications and reduces the cost of such procedures.415Shrestha et al. [4] studied the effect of part size and geometry on the thermal history and mechanical properties of laser beam powder bed fusion (LB-PBF) 17-4 PH stainless steel (SS). They showed that small changes in the size and geometry of the part will change the thermal history experienced by the part, thereby influencing the defect size and distribution. They also reported better fatigue resistance for specimens machined out from square bars as compared to the ones fabricated slightly oversized than the near net-shape.

They also reported better fatigue resistance for specimens machined out from square bars as compared to the ones fabricated slightly oversized than the near net-shape. This was attributed to the higher bulk temperature in former specimens and led to the formation of lower porosity level.Gu et al. [7] studied the effects of energy density on the microstructure and porosity formation in LB-PBF 17-4 PH SS by altering the scan speed and laser power. Results showed that lower porosity level will be achieved by adjusting the energy density in the range of 44-81 J/mm3. However, they showed that energy density cannot be a good indication of the porosity level. For instance, increased scan speed and laser power led to the lower energy density in the range, however, the porosity level increased due to the balling phenomenon. Sun et al. [8] produced a unique crystallographic lamellar microstructure (CLM) in LB-PBF 316L SS via controlling the energy density.

Sun et al. [8] produced a unique crystallographic lamellar microstructure (CLM) in LB-PBF 316L SS via controlling the energy density. The newly developed texture for this material comprised of two grain orientations, <001> and <011>, which enhanced the mechanical properties significantly.Thermal history was also observed to be sensitive to the build platform temperature during fabrication through AM process. It was reported that preheating the build platform can lead to lower residual stress, more homogenized microstructure, and consequently, better mechanical properties (e.g. ductility, fatigue resistance, and etc.) as compared to the condition without preheating the build platform [9]. Yang et al. [10] investigated the effect of preheating the build platform at constant energy densities on the microstructure of Al-Mg (-Sc)-Zr. They observed columnar to equiaxed grain morphology as the build platform preheated at 200 °C.

They observed columnar to equiaxed grain morphology as the build platform preheated at 200 °C. They attributed this microstructure evolution to the reduced thermal gradients and increased remelting zone volumes. Liu et al. [11] investigated the effects of various build platform preheating temperatures on the texture evolution of TiAl alloy via LB-PBF process. They reported significant differences between the grain orientations obtained under different preheating temperatures.Accordingly, due to the complexity of changing process parameters such as laser power, scan speed, scan strategy, etc., fabricating parts by preheating the build platform can be a solution to manufacture parts with more homogeneous microstructure and less residual stress. Austenitic stainless steel 316L is one of the most attractive materials in various industrial sectors. Moreover, its biocompatibility and corrosion resistance make this material suitable for medical applications such as bone implants [12].

Moreover, its biocompatibility and corrosion resistance make this material suitable for medical applications such as bone implants [12]. The goal of this study is to investigate the effects of build platform preheating on the microstructure and porosity level of LB-PBF 316L SS. Numerical simulations are carried out using ANSYS Additive Manufacturing module to calculate the variation in cooling rates. Numerical simulations are then validated by experimental observations of the melt pool dimensions and microstructural variation for different cooling rates. Eventually, the porosity size and distribution are evaluated for specimens fabricated with preheating the build plate and compared with the ones fabricated without preheating the build plate. Tensile deformation behavior is assessed for each condition and correlated to the microstructure and porosity size and distribution.416Experimental ProceduresIn this study, 316L stainless steel (SS) powder provided by LPW Inc.

with the chemical composition listed in Table 1 was used to fabricate specimens. The specimens were fabricated via a laser beam powder bed fusion (LB-PBF) method in the vertical direction using Renishaw AM 250 machine.Table 1. Chemical composition of 316L SS powder, provided by LPW Technology Inc.CrNi MnSi Nb MoNOPSFe(Wt. %) 16-18 4.03 0.20 0.10 0.33 2.0-3.0 <0.10 <0.10 <0.04 <0.03 Bal.Two sets of specimens were fabricated under argon shielding gas, without and with preheating the build platform to 150°C. Specimen geometry is shown in Figure 1. For better tracking each set of specimens in the following context, specimens fabricated without preheating the build platform, and the ones fabricated with preheating the build platform to 150 °C are designated by NP and P150, respectively.rcl 120 ~ l-1 1· :1tl201IB;;/ .Figure 1.

Drawing of sub-sized tensile specimens based on the ASTM E8 standard.Thermal properties were predicted for each condition (i.e. NP and P150) employing the commercial ANSYS Additive Manufacturing module. In this module, deposition thickness, scan speed, hatch spacing, layer time interval, preheating temperature, and powder and gas convection coefficients were input to the software. Software input parameters were selected based on the manufacturing machine (i.e. Renishaw AM 250) and optimized process parameters for 316L SS (i.e. layer thickness of 50 (cid:541)m, hatch spacing of 0.11 mm, laser power of 200 W, and a scan speed of 1833 mm/s). The layer time interval in the gauge section was 65 s recorded during the fabrication process. Material (i.e. 316L SS) was selected from the ANSYS database with pre- defined temperature-dependent properties including density, thermal conductivity, and heat capacity.

Material (i.e. 316L SS) was selected from the ANSYS database with pre- defined temperature-dependent properties including density, thermal conductivity, and heat capacity. Computer-aided designed (CAD) or stereolithography (STL) files of fabricated specimens were imported directly to ANSYS to perform the numerical analyses. In addition, the build platform was considered in the design in order to consider the heat dissipation to the build plate. The software outputs were the maximum and minimum temperatures in each layer depending on the fabrication time. In ANSYS Additive Manufacturing module, temperatures are obtained by solving the partial differential equation (PDE) of conduction heat transfer given in Eq.

The (cid:2025), (cid:1863), and (cid:1855) are material temperature-dependent density, thermal conductivity, and heat capacity, respectively. In addition, (cid:1869)(cid:4662) is the heat input per volume due to laser beam source, which is considered via a Gaussian distribution [13]. Convection and radiation heat losses can be also considered as the boundary conditions to solve the PDE [13]. However, in this study, the effects of convection and radiation were neglected to simplify the model. Finally, specimen drawings were cut in the middle of the gauge and used for numerical analyses in order to focus on the thermal history of the gauge section as well as decreasing the processing time. The heat transfer coefficients of powder and shielding gas (i.e. argon) were selected as a relatively small number to ensure that conduction is the main source of heat dissipation.

The heat transfer coefficients of powder and shielding gas (i.e. argon) were selected as a relatively small number to ensure that conduction is the main source of heat dissipation. Two simulations were performed based on the two experiments, one with preheating the build platform of 150 °C and one without preheating. In addition, the mesh size was selected as 0.2 mm to ensure the mesh discrepancy of the results. The build plate was meshed with an element size of 5 mm since it was kept at a constant temperature during the fabrication time.For microstructure characterization, some specimens were fabricated to the half-height of the regular ones. Those specimens were transversely cut perpendicular to the final laser tracks to look into the melt pool’s geometry. Microstructure samples were then ground using 320-2500 grit sandpapers to the mirror-like surface finish. Prior to reveal the microstructure, porosity measurement was carried out using KEYENCE digital optical microscope.

Prior to reveal the microstructure, porosity measurement was carried out using KEYENCE digital optical microscope. Specimens were further etched to reveal the microstructure, showing the grains and melt pool dimensions on the most top fabricated layer.Tensile tests at 0.001 s-1 strain rate were conducted at room temperature using an MTS Landmark servohydraulic test system with 100 kN load cells. Tests were conducted in two steps, strain-controlled mode followed by displacement-control mode. Strain-control mode step was performed up to 0.045 strain using an MTS extensometer to record the strain. However, due to the limited displacement of extensometer and to avoid any damage, tests were paused to remove the extensometer. Subsequently, tests were continued in displacement-controlled mode to fracture with a velocity equivalent to the same strain rate. To obtain the elongation to failure, gage sections were marked prior testing, and measured after tests were done.

To obtain the elongation to failure, gage sections were marked prior testing, and measured after tests were done. Two tests were conducted for each condition to observe the consistency of the results. To understand the fracture behavior in each condition, fractography analysis was carried out for all the tested specimens using KEYENCE digital microscope.Results and DiscussionThermal Simulation and MicrostructureFigure 2 shows a time window of the interlayer temperature fluctuation during fabrication for each condition (i.e. P150 and NP). In simulations, the top layers of parts were fabricated at the time 31305 s, when the laser scanned for the last time. It can be observed that in every deposited layer, the temperature reaches to the melting point temperature and subsequently drops from that temperature to the desired temperature (i.e. 150 °C for P150 and 22 °C for NP).

150 °C for P150 and 22 °C for NP). It can be observed418that temperature decreases faster from 1370 °C to 22 °C for the NP process while it decreases from 1370 °C to 150 °C at the same time for P150 process. This trend shows that NP condition possesses a higher cooling rate than P150 condition.1600...,_Preheated1400~ -Non-preheatedMelting tempe.-atu.-e line120010008006004002000 31000\ \ \ I 0-----~ 31100®-----~310503115031200 Time (s)3125031300\ \ \ \ \ \ ~3135031400Figure 2. Temporal distribution of maximum temperature along the fabrication direction (z) for both P150 and NP conditions. Middle of the gauge section is scanned at time 31,305 s.Masoomi et al. [14] reported that cooling rates are governed by the conduction, convection, and radiation within the heat affected zone (HAZ), i.e., the solid region in the vicinity of the melt pool, shielding gas, and the walls of the chamber, respectively.

Since the effects of convection and radiation were neglected in this study, heat is dissipated by conduction in the HAZ. The conduction heat dissipation is dependent on the thermal gradient between two points, thereby, cooling rates are directly related to the temperature differences between these two points, which is in accordance to Newton law of cooling. Therefore, a higher temperature difference between two points can derive a higher cooling rate, which can be also observed in Figure 2. The cooling rate values were obtained from the simulation to be 1.43×106 °C/s for P150, and 1.65×106 °C/s for the NP condition. Some studies have shown that lower thermal gradients, and consequently, lower cooling rates, can result in melt pools with larger dimensions [15].

Some studies have shown that lower thermal gradients, and consequently, lower cooling rates, can result in melt pools with larger dimensions [15]. Therefore, larger melt pool dimensions including melt pool depth can be expected for the parts that are preheated.Figure 3(a) and (b) show the melt pool dimensions for specimens fabricated without preheating (NP), and with preheating (P150) the build platform, respectively. The average width to depth ratios (i.e. Wavg /davg) of the melt pool is calculated for NP and P150 conditions to be 1.38 and 1.7, respectively. It can be seen that P150 condition has larger melt pools as compared to the NP one, which is consistent with the lower cooling rate calculated for the former condition (i.e.419P150).

It can be seen that P150 condition has larger melt pools as compared to the NP one, which is consistent with the lower cooling rate calculated for the former condition (i.e.419P150). Moreover, melt pool size affects the porosity level due to the gas entrapment as argon cannot be dissolved in steels.Comparing the microstructure obtained for P150 with NP condition, the cellular structure is seen in the former condition (shown with white arrows in Figure 3(b)). Based on the thermal simulation results shown in Figure 2, the decrease in the cooling rate by preheating the build platform is due to the lower temperature gradient induced between the layer melted and the previous fused layers. This results in not only grain structure modification but also the texture evolution [11]. Typically, formation of fine cellular-dendritic microstructure has been reported as the common microstructure produced in the AM process [16,17]. This is due to the heat dissipation toward the build platform.

This is due to the heat dissipation toward the build platform. Formation of the cellular-dendritic structure has been also reported for LB-PBF 316L SS [18]. However, it has been shown that preheating the build platform results in more equiaxed grains [10]. Moreover, it has been reported that increasing the laser power (i.e. higher energy density) leads to the formation of equiaxed cellular structure in 316L SS [19]. Accordingly, the results obtained found the preheated platform condition to be consistent with the literature.Figure 4(a) and (b) show the transverse plane in the gage section polished to characterize the porosity size and distribution for NP and P150 conditions, respectively. In addition, the porosity size and distribution histograms for each condition are plotted. Comparing the polished sections, the porosity level decreased for the specimens fabricated with preheating the build platform.

Comparing the polished sections, the porosity level decreased for the specimens fabricated with preheating the build platform. The statistics results of porosity show that frequency of pores in P150 condition is higher in low size classes (i.e. [11-18] (cid:645)m) as compared to the NP condition. Consequently, fewer large pores are seen for P150 condition. A larger and higher number of pores in NP condition is related to the relatively higher cooling rates in this condition as compared to P150. In fact, the entrapped gas does not get the chance to escape from melt pools due to rapid solidification, thereby, it results in more porosities.(a)(b)Figure 3.

In fact, the entrapped gas does not get the chance to escape from melt pools due to rapid solidification, thereby, it results in more porosities.(a)(b)Figure 3. Melt pool characterization on the most top layer: (a) non-preheated, and (b) preheated build platform.420(a)(b)NPPl50606050504040c g 30 uc g 30 u2020101000[11-18)(18-25[(25-32[(32-39)(39-46)(46-53)(53-60[Jll -1 8)(18-25 [(25-32)(32-39[(39-46[(46-53[Pore Size (µm)Po.-c Size (µm)Figure 4. Porosity measurement results in the gage section: (a) polished section and porosity size distribution for non-preheated, and (b) polished section and porosity size distribution for preheated build platform conditions. Note the size classes for each condition.Tensile propertiesMonotonic tensile deformation behavior of specimens fabricated with and without preheating the build platform is shown in Figure 5.

Note the size classes for each condition.Tensile propertiesMonotonic tensile deformation behavior of specimens fabricated with and without preheating the build platform is shown in Figure 5. The strain-controlled mode portion of the tensile test is shown in Figure 5(a), and the stress-displacement curve from the beginning of the tests to the fracture is shown in Figure 5(b). Moreover, the tensile results are compared with [20] and wrought material [21–23], listed in Table 2. The ultimate tensile strength and elongation to failure (El%) of the present alloy increased as compared to those of additively manufactured and wrought counterparts. Interestingly, preheating the build plate increased the El% and yield strength ((cid:305)y) of the material by approximately 14% and 4%, respectively. This is attributed to the microstructure and texture evolution. As seen in Figure 3(b) shown by white arrows, the cellular structure has been developed in the P150 condition.

This is attributed to the microstructure and texture evolution. As seen in Figure 3(b) shown by white arrows, the cellular structure has been developed in the P150 condition. These cellular structures, either dendritic- or equiaxed-cellular structures, are made of high dislocation density as a result of the high cooling rate during fabrication [18]. However, in P150 condition, dislocation rearrangement and recovery may occur and lead into the equiaxed-cellular structure, as the previously fused layers are still at high temperature [24]. Accordingly, dislocation density is decreased, and the microstructure might be considered similar to the annealed condition. Therefore, the material is capable of421(53-60[accommodating higher plastic deformation by generation of geometry necessary dislocations (GND) during straining the material [25].700700600500;.- 400 ~ "' 300 "' ~ rii 200100--/,·__ ......................... ~ ..... -..... ~ ......

Monotonic tensile properties of 316L SS fabricated under argon shielding gas in both non-preheated and preheated build platform conditions. Build platform conditions Non-preheated Preheated LB-PBF 316L [20] Wrought [21][22,23](cid:305)y (MPa)(cid:305)u (MPa)El%448 465617 61872 8248559458255-310535-62330-40FractographyFigure 6 shows the fracture surfaces of tensile specimens for the non-preheated and preheated conditions. As seen, both fracture surfaces exhibited cup and cone type failure and consisted of fibrous and shear region. These fracture features are consistent with the tensile results obtained, showing ductile behavior of the material. As seen in Figure 6(a), NP specimen has larger fibrous region than that of P150 specimen, however, tensile results show lower ductility for the former condition. This is due to the lower structural integrity of NP condition resulted from higher porosity level as compared to P150 condition.

This is due to the lower structural integrity of NP condition resulted from higher porosity level as compared to P150 condition. Moreover, although dimples exist in both conditions (shown by red arrows), the specimens fabricated on the preheated build platform consist of deeper dimples (~ 25-27 (cid:541)(cid:80)) as compared to those of the ones fabricated on the non-preheated build platform (~ 15-21 (cid:541)(cid:80)). Deeper dimples are the sign of higher ductility in the material, which can be seen that the P150 specimens possess higher ductility than NP ones. Shallower dimples in NP condition can also be due to less structural integrity in this condition due to the presence of more42214porosities and lack- of-fusion (LOF) defects as compared to the P150 condition (see Figure 4). Shrestha et al. [20] also showed various LOF defects in the fracture surfaces of LB-PBF 316L SS fabricated with non-preheated build platform condition.(a)(b)Figure 6.

Shrestha et al. [20] also showed various LOF defects in the fracture surfaces of LB-PBF 316L SS fabricated with non-preheated build platform condition.(a)(b)Figure 6. Tensile fracture surfaces of LB-PBF 316L SS specimens: (a) non-preheated, and (b) preheated build platform conditions.Conclusions and future studyThe effect of preheating the build platform on the microstructure and mechanical properties of LB- PBF 316L SS was investigated. The following conclusions can be drawn based on the results:1. The cooling rate of P150 condition was decreased as compared to that of the NP condition. This resulted into formation of equiaxed cellular structure.2. Preheating the build platform to 150 °C increased the ductility of material by 14%. This is attributed to the more homogenized microstructure as well as cellular structure with geometry necessary dislocations (GNDs), thereby material accommodated more plastic strains during deformation.3.

The porosity size and distribution were decreased by preheating the build platform. The critical defect size for the NP condition was in the range of [32-39], while it was in the range of [17-24] for the P150 condition.4. Tensile fracture surfaces of P150 specimens were found to have less fibrous region as compared to NP ones. However, due to the higher structural integrity of P150 (i.e. less porosity and lack of fusion defects) than NP specimens, ductility increased without any change in the material strength.Due to the significant effect of preheating the build platform on reducing the porosity size and distribution, fatigue performance of the material assumed to be enhanced as compared to the condition without preheating the build platform.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceThe Effect of Print Speed on Surface Roughness and Density Uniformity of Parts Produced Using Binder Jet 3D PrintingK. Myers1, A. Paterson1, T. Iizuka2, and A. Klein11The ExOne Company, 127 Industry Boulevard, North Huntingdon, PA, 156422ExOne KK, 161-5 Haneo, Odawara-shi, Kanagawa Japan, 256-0804AbstractOne of the main benefits of binder jetting is the ability to print quickly compared to other metal additive manufacturing methods. Demand for higher throughput continues to increase, but the effects of faster print speeds on part outcomes are not yet clearly understood. MIM powders are used to achieve optimal density and surface finish. Printing at slower speeds results in densities near 98% and average surface roughness values as low as 4 μm (Ra), in the as-sintered condition.

Printing at slower speeds results in densities near 98% and average surface roughness values as low as 4 μm (Ra), in the as-sintered condition. In this study, spread speeds were varied in order to understand the effect of print speed on surface roughness. 316L D90 -22 μm powder was used to print with 3 different spread speeds, 2 different layer thicknesses, and 2 different printhead droplet sizes. The surface finish and density were quantified for the sintered parts that were oriented at 0, 22.5, and 45 degrees with respect to the Z- direction.IntroductionBinder jet additive manufacturing is a powder-based technique that uses an inkjet printhead to selectively deposit binder to form a desired 3D shape. Binder jetting closely resembles powder manufacturing and injection molding techniques in the sense that a green body is formed with a combination of binder and powder, which subsequently undergoes a debinding and sintering process to create a fully dense final part.

Compared to laser or electron beam metal additive manufacturing techniques, binder jetting has the highest potential to be competitive for large volume production. This is due to relatively quick printing times and the ability to post-process many parts in large batch furnaces or continuous feeding furnaces, in much the same way as metal injection molding (MIM). The time required to print a part via binder jetting is the sum of four key processes: powder dispensing, powder spreading, binder dispensing, and binder drying. While the layer thickness of the print can also affect printing time, thickness is typically determined by a relationship of approximately 2x the largest particle size in the distribution [1].Past work has documented the difficulties of printing with powders smaller than 5 (cid:645)m while using older spreading and dispensing techniques.

Most commonly, large porosity was found in parts due to improper spreading of powder, resulting in lower overall mechanical strength for the part [2]. This is commonly caused by van der Waals cohesion forces in fine powders [3]. The printing time (the time it takes to dispense binder droplets) can be decreased by either increasing the number of printing modules in the machine itself, or by increasing the speed at which the printhead traverses across the bed. However, increasing the printhead traverse speed can122negatively affect the feature resolution and dimensional accuracy [4]. For powder spreading, the amount of powder that is dispensed and used for spreading directly affects the green density and uniformity of the printed parts. For example, higher amounts of powder in front of the roller can result in higher green densities [5-6]. Drying time is a function of both the amount of binder (saturation) that is deposited each layer as well as the binder’s composition.

Drying time is a function of both the amount of binder (saturation) that is deposited each layer as well as the binder’s composition. Not allowing enough time for binder dispersion and drying can lead to lower green body strengths [7]. Additionally, not allowing enough time for liquid component of the binder to evaporate can lead to smearing of the part, creating rougher surface finishes and dimensional inaccuracy.There are only a few literature articles on the impact of spreading speed on green density and powder packing consistency. Studies have found that slower spreading speeds have produced the highest green densities [7-8]. In both studies, a spread speed of 6 or 10 mm/sec was the highest speed chosen, respectively. In other studies, Mostafaei et al. used a spread speed of 15 mm/sec to print Inconel 625 [9], and Nandwana et al. reports that using a spread speed over 4 mm/sec when printing Inconel 718 would lead to non-uniform spreading and delamination [10].

reports that using a spread speed over 4 mm/sec when printing Inconel 718 would lead to non-uniform spreading and delamination [10]. Typically, the powder spreading speed has the largest impact on determining the time to print one layer, with a typical spreading speed of 5 mm/sec or slower for particle size distributions similar to -22 μm. The average layer time for the ExOne Innovent+ system is 75 seconds when spreading powder at 3 mm/sec. With this set of parameters, 29% of the layer time is from spreading powder. That percentage is much higher when using larger build envelopes, like the X1 25Pro and MFlex.In a production environment, consistency is crucial to reliably making parts with close tolerances. Such factors as machine wear, machine cleanliness, the number of times powder is recycled, and humidity in the air can alter the flow rate and flow consistency out of the hopper during powder dispensing.

Being able to precisely control the amount of powder that is being spread is the key to consistency in green body density. Figure 1 is a schematic of ExOne’s Advanced Compaction Technology (ACTTM), developed by ExOne (patent pending), for powder spreading, where LT is the layer thickness, RT is the roughing powder thickness, CR is the compaction ratio, (cid:2033)(cid:3020) is the rotational speed of the smooth roller, (cid:2033)(cid:3019) is the rotational speed of the roughing roller, and (cid:646) is the spread speed. As mentioned before, the counter rotating roller mechanism of spreading powder has been shown to result in packing densities that are dependent on the amount of powder in front of roller. The ACT works by using a roughing roller ahead of a smooth roller that removes or “cuts” the loose powder layer to a known height and packing density, and has minimal effect on compaction.

The ACT works by using a roughing roller ahead of a smooth roller that removes or “cuts” the loose powder layer to a known height and packing density, and has minimal effect on compaction. Preliminary results have shown that the roughing nature of the roller allows the packing density of the powder to be unaffected by the mass of powder that is in front of the roller. This demonstrates that the powder layer created by the roughing roller will be consistent, as long as enough powder is dispensed to the height of LT + RT.123.. .,,,.----- ...

This demonstrates that the powder layer created by the roughing roller will be consistent, as long as enough powder is dispensed to the height of LT + RT.123.. .,,,.----- ... Roughing Roller,,......,,;' ' '/ICR= RT LTI\Smooth Roller w,I I IW R\ \I ~I I I I\ \ \I I~I' ' ' ...,,/I I I...;Excess Powder...1--c _..,,,,.,,, - I ILoose PowderRT-------~-Figure 1 – Schematic of the advanced compaction technology used during the powder spreading process.This research was performed in order to investigate how the powder spread speed using the ACT affects the green body density as well as the surface roughness of the sintered part. The impact of layer thickness and droplet size on surface roughness was also investigated.ExperimentalA 90% -22 μm 316L powder was used during this testing. The particle size distribution was determined to be a d90 of 15.6 μm, a d50 of 7.6 μm, and a d10 of 3.5 μm.

The particle size distribution was determined to be a d90 of 15.6 μm, a d50 of 7.6 μm, and a d10 of 3.5 μm. The particle size was determined by a Shimadzu SALD-3001 laser diffraction particle size analyzer. All parts were printed on an ExOne Innovent+ binder jet printer using the ExOne solvent binder at 70% saturation, a smooth roller rotational speed and a roughing roller rotational speed were held constant, and a compaction ratio of 2. The layer thickness and droplet size were tested with the combination of either a 50 μm layer thickness with a 30 pL droplet or a 30 μm layer thickness with a 10 pL droplet. The roller height difference for the Advanced Compaction Technology was set at 100 μm. 15 coupons (25.4 x 25.4 x 6.35 mm) were printed in each trial. Of those, 5 coupons each were printed at angles of 0, 22.5, and 45 degrees in relationship to the Z-direction (see Figure 2). After printing, all samples were cured for 8 hours at 200°C in air.

Of those, 5 coupons each were printed at angles of 0, 22.5, and 45 degrees in relationship to the Z-direction (see Figure 2). After printing, all samples were cured for 8 hours at 200°C in air. Following the curing stage, the coupons were depowdered and physically measured with calipers and weighed to obtain density measurements.124Figure 2 – The build layout of each trial as shown in the Innovent+ software.All coupons were sintered in a graphite hot zone furnace with partial pressure hydrogen to a temperature of 1380°C. After sintering, the coupons were once again physically measured with calipers and weighed for density calculations. The surface roughness of each coupon was measured 5 times with a Phase II portable surface roughness gauge, which has a diamond stylus that is accurate to 0.001 μm.

The surface roughness of each coupon was measured 5 times with a Phase II portable surface roughness gauge, which has a diamond stylus that is accurate to 0.001 μm. A length of 1 mm was sampled in each test in accordance with Figure 3.125a)b)Testing DirectionLayer ThicknessFigure 3 – a) Schematic of a coupon being tested 5 times in various positions across the surface of the coupon and b) schematic of the stylus traverse direction against layer thickness of angled coupons.3) S htifThe surface roughness of one coupon from each of the printing conditions was also tested with a Keyence VX-6000 light microscope in a 3D plot of the topology. The 3D topology scan was performed across a 12.5 x 2.5 mm plane on the top surface of the coupons. The microscope scanned a rate of 0.3 mm/sec at 500x magnification.Results and DiscussionThe green and sintered density values from all six printing trials can be seen in Table 1.

The microscope scanned a rate of 0.3 mm/sec at 500x magnification.Results and DiscussionThe green and sintered density values from all six printing trials can be seen in Table 1. Table 1 shows that in all trials, when the powder spread speed is increased, the standard deviation of the density increases. This demonstrates that the slowest spreading speed allows for the most consistent powder packing density. The average green density decreased with the larger layer thickness of 50 μm, but stays relatively consistent with increased powder spreading speed. Similar to MIM, a higher degree of consistency in green body density creates consistent and repeatable shrinkage during sintering.

Similar to MIM, a higher degree of consistency in green body density creates consistent and repeatable shrinkage during sintering. Although the coupons with the 50 μm layer thickness have a lower average green density than the 30 μm layer thickness, the coupons were able to sinter to over 98% dense.When evaluating the effect coupon angle has on green density, the average density decreases with increased coupon angle (see Table 2). This is consistent with literature which has found that porosity in the green state of binder jetted parts is mostly found between the conjoining layer thicknesses due to particle ejection from the impact of the droplets [11]. This porosity effect can also manifest in the sintered stage, showing lines of porosity in the Z-direction [12]. The porosity increases in the angled coupons due to the increasing the number of layers needed to print these parts.

Dev.% 0.11 0.28 0.25 1.34 0.41 0.17 0.42 0.37 0.30 0.22 0.23 0.54 1.77 1.18 1.13 0.78 2.82 0.29127Green Density vs Coupon AngleTrial 1■59.00Trial 2■Trial 3■)57.00Trial 4■%( y t i s n e D n e e r G55.0053.0051.00Trial 5■Trial 6■49.0047.00022.5 Coupon Angle (°)45Figure 4 – The average green density profile for each Trial in respect to coupon print angle.The surface finish data from the Phase II surface roughness gauge can be seen in Table 3 and are plotted in Figures 5 and 6. The data shows that in the case of both the 10 pL and 30 pL printhead trials, the surface roughness (Ra) increases with both spread speed and coupon angle. The averages of all of the trials showed that the lowest surface roughness (Ra) was 4.05 (cid:645)m, and was found on the sample printed at 0° from the horizontal with a 30 pL printhead and a 50 μm layer thickness.

The highest Ra value was 9.78 μm, which was printed with a 30 pL printhead and 50 μm layer thickness. The data shows that Ra value approaches a plateau between 50 and 125 mm/sec spread speed. The data from the Keyence microscope is in agreement with the data found from the Phase II roughness gauge, showing the same trends for increased surface roughness with spread speed and angle. The plots of the surface roughness found with the Keyence microscope can be seen in Figures 7 and 8. The surface roughness values are also in agreement with literature values of sintered parts with similar particle size distributions [9,13]. A topology map of the 10 pL sample printed at 0 degrees and 3 mm/sec taken with the Keyence microscope at 100x magnification can be seen in Figure 9.

A topology map of the 10 pL sample printed at 0 degrees and 3 mm/sec taken with the Keyence microscope at 100x magnification can be seen in Figure 9. The topology map shows uniform surface with typical hills and valleys, with no unique surface topology defects or spikes.128Table 3 – The average coupon surface roughness values of each trial, taken from the Phase II analyzer.Trial# 1 1 1 2 2 2 3 3 3 4 4 4 5 5 5 6 6 6Spread Speed mm/sec 3 3 3 50 50 50 125 125 125 3 3 3 50 50 50 125 125 125Printhead Angle AveragepL 30 30 30 30 30 30 30 30 30 10 10 10 10 10 10 10 10 10° 0 22.5 45 0 22.5 45 0 22.5 45 0 22.5 45 0 22.5 45 0 22.5 45Ra μm 4.05 7.66 7.27 5.43 9.50 9.42 6.40 9.78 9.20 5.02 6.70 7.27 5.58 7.86 7.86 5.02 7.21 7.87St. Dev.

The relationships were characterized for the combination of a 50 μm layer thickness and 30 pL printhead, and a 30 μm layer thickness and 10 pL printhead. The results of the testing are the following:The average green body density decreases and the standard deviation increases with increasing spread speeds.

The average density did not decrease with spread speed with the 30 μm layer thickness and 10 pL printhead, but the standard deviation increased.The average green body density was slightly higher with the 30 μm layer thickness (54%), than the 50 μm layer thickness (50-52%).The average green body density decreases slightly with increasing coupon orientation angle in relationship to the Z-direction.Although there is a deviation in green body densities from various process settings, all coupons sintered to an average green body density of 97.5% theoretical or higher.The surface roughness (Ra) typically increases with faster spreading speeds and increased coupon orientation angle.132The smallest average surface roughness value (Ra) was found to be 4.05 μm with the 50 μm layer thickness and 30 pL printhead on a flat surface with a spread speed of 3 mm/sec.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of Printing Parameters on the Internal Geometry of Products Manufactured by Fused Filament Fabrication (FFF)Benjamín A. Moreno-Núñez†, *, Cecilia D. Treviño-Quintanilla, *, Juan Carlos Espinoza-García†, Esmeralda Uribe-Lam† and Enrique Cuan-Urquizo.†School of Engineering and Sciences, Tecnologico de Monterrey, QRO 76130 Institute of Advanced Materials for Sustainable Manufacturing, Tecnologico de Monterrey, QRO 76130 *Corresponding authorsAbstractThe internal geometry of a 3D-printed product determines its mechanical properties. In Fused Filament Fabrication (FFF) the filaments that build up the internal geometry suffer from variations that have not been sufficiently studied.

In Fused Filament Fabrication (FFF) the filaments that build up the internal geometry suffer from variations that have not been sufficiently studied. This research focused on identifying the parameters that most affect the filaments and finding the optimum values to reduce their variations. A fractional factorial design of experiments was used to detect the printing parameters of FFF that most affect the width of extruded filaments, these results were also statistically analyzed. A response optimization was done to obtain the values of the printing parameters that will give the closest width of extruded filaments to the nozzle of the 3D printer used.

A response optimization was done to obtain the values of the printing parameters that will give the closest width of extruded filaments to the nozzle of the 3D printer used. Results showed layer height has the largest impact on filament width variation.Keywords: additive manufacturing, fused filament fabrication, statistical analysis,fractional factorial design, internal geometry, thermoplastics.IntroductionA modern manufacturing technique used to create structural, athletics, medical, and aeronautical devices is additive manufacturing (AM) [1–4]. AM can contribute to reduce production costs in low-volume lines or industries involved in new product development [2]. A process that is part of AM processes is Fused Filament Fabrication (FFF), also called Fused Deposition Modelling (FDM), a technology first developed by Stratasys® [5] to accomplish 3D- printed products.

The FFF process consists in heating up and extruding a thermoplastic filament through a nozzle that deposits it on specific zones and uses a layer-by-layer method upon finishing the desired geometry.During the FFF process, the G-code is the primary component and is in charge of modifying the printing parameters [6–8]. Some of the printing parameters that G-code can control are nozzle movement, extrusion temperature, feed rate, bed temperature, printing speed, fan, or layer thickness [9]. The printing parameters have a direct effect in the thermoplastics used in FFF, they can create defects, like weak bonding or porosity [10] on the structure due to the interaction of mechanical operation and thermal changes [11,12].

These structural defects will impact the mechanical response, diminishing mechanical properties on products made with thermoplastic filaments [11,13–17].The dimensional accuracy of FFF products is a critical point for the mechanical response of the FFF products, this accuracy can be affected by the printing parameter variations. The FFF products have lower dimensional accuracy (2.2% of volumetric variation) when compared to1832foamed or injected thermoplastic products (1.3% of volumetric variation) [18]. Dimensional accuracy can be affected by infill density, extrusion rate, feed rate, nozzle diameter, the length of the extruder travel, layer accumulation, or by the tangential velocity of the printing path [19–22].One part of dimensional accuracy is the filament width, which can affect the mechanical response of FFF products because it is related to the infill geometry of them.

The infill geometries are included in FFF products to have high strength and a relative low mass [23]. A variation in the infill pattern, layer thickness, or lack of uniformity will cause variating strength, quality, and behavior of FFF products [24,25]. It has been found that the filament can be affected by nozzle speed, the higher the nozzle speed, the smaller the filament diameter [26]. Also, the viscoelastic stresses produced in the 3D-printing process can affect the extruded filament dimensions [27].Statistics have been used in FFF to analyze physical and mechanical responses, the most common methods used are response surface [28–31], Taguchi methodology [32,33], central composite design [31,34,35], and fractional factorial design [29,36]. These models have been used in the FFF process to make regression models that can help to optimize the printing process parameters and improve the properties of the final 3D-product [20,34,36–39].

These models have been used in the FFF process to make regression models that can help to optimize the printing process parameters and improve the properties of the final 3D-product [20,34,36–39]. Regression analyses like surface methodology, support vector regression, or multivariable linear regression have been used in FFF processes and products to make regression models that predict and optimize flexural properties [30], dynamic mechanical performance [39], tensile strength [40–42], creep rate [43], flexural properties [44], warpage [45,46], or surface roughness [28,47,48].

More modern techniques like artificial neural networks [22] or Gaussian process regression [48] can give more accurate results but require more factors that can be translated to time consumption [22].Full factorial design has been used to analyze the physical responses of FFF products, for example Vyavahare et al., they used a factorial design to evaluate the effect of five parameters on surface roughness, dimensional accuracy, and time of fabrication on ABS 3D-printed products [28]. Sood et al. used a factorial design to analyze the sliding wear of FFF products, using the response of five parameters adjusted to a factorial design and train an artificial neural network [31]. Santana et al. used a Taguchi L8 orthogonal array and a full factorial design to study the effect of 8 parameters on the dimensional quality of 3D-printed PLA products [20]. Armillotta et al. analyzed the warpage effect of ABS 3D-printed products using 4 printing parameters and a full factorial design [46].

Armillotta et al. analyzed the warpage effect of ABS 3D-printed products using 4 printing parameters and a full factorial design [46]. García Plaza et al. used a full factorial design to study 3D-printed PLA geometric properties and the effect on them of 3 printing process parameters, also using the obtained data to feed an artificial neuronal network and make predictions [22].Some authors have studied the extrusion process and the impact on filament width. Gleadall et al. made a model that simulates the extrusion process of FFF and reached an approximation of filament width with a 5% of error in the 3D-generated-models of engineering scaffolds [49]. Hebda et al. made a model that can predict the filament width, considering the deposition height equal to the nozzle set gap height [50]. Yousefi et al. analyzed the extrudate swell of polycaprolactone (PCL) scaffolds made on a 3D-bioplotter, using an I-optimal, split-plot experiment.

Yousefi et al. analyzed the extrudate swell of polycaprolactone (PCL) scaffolds made on a 3D-bioplotter, using an I-optimal, split-plot experiment. In this experiment it has been analyzed the effect of needle diameter, temperature, pressure, and dispensing speed into swell, density, and modulus of the 3D-printed samples. With the results of the experiment, they made 3 regression models for the 3 responses, having percentage differences of 9.5% for swell, 19% for density, and 29% for modulus [51].The products obtained from FFF suffer from dimensional inaccuracy and do not meet the requirements to be used as functional parts because of the high risk of failure due to variations in1833the structure [52]. The aim of this research is to diminish the variation in the width of the individual filaments used to create the infill geometry found inside the products made by FFF, resulting in a homogeneous structure that can make functional 3D-printed products.

To achieve this, there were selected the most common FFF printing parameters that can be modified in commercial slicer software, to obtain with their modifications the lower error in their width. Also, this research aims to obtaining the printing parameter values that can reduce the mentioned variation of filament width by obtain regression models that can be set depending on the materials and type of 3D printer used.Materials and MethodsMaterialsThe samples were fabricated by FFF, using three different thermoplastic filaments, Polylactic Acid (PLA), Acrylonitrile Butadiene Styrene (ABS), and Polycarbonate (PC). The filaments have a diameter of 1.75 mm and 2.85 mm, provided by Colorplus 3D (Mexico) and XinYuDa (XYD) Plastic Products Factory (China).

The filaments have a diameter of 1.75 mm and 2.85 mm, provided by Colorplus 3D (Mexico) and XinYuDa (XYD) Plastic Products Factory (China). All filaments were selected white to diminish the dimensional variation caused by the color of the filament [53].3D Printing ProcessTo fabricate the samples used in this research, two different printer types were used, a Cartesian printer, Ultimaker 2+ (Netherlands), and a Delta printer, RostockMax V3.2 (USA). The nozzles used to obtain the parts have a diameter of 400 𝜇𝑚, and 500 𝜇𝑚, for Cartesian and Delta printers, respectively. The samples were configured to have 4 alternated layers, 2 with horizontal printing direction and 2 with vertical printing direction, as presented in Figure 1A, to accomplish a 2D square lattice infill geometry. The separation between individual filaments was 1.5 mm, as presented in Figure 1B.Figure 1. A. Scheme of fabricated samples and B.

The separation between individual filaments was 1.5 mm, as presented in Figure 1B.Figure 1. A. Scheme of fabricated samples and B. Separation between individual filaments.To analyze the effect of the parameters on the extruded filament width, a fractional factorial design of experiments was made for each material and 3D-printer used in this research. Five printing parameters were selected to study their effect on extruded filament width, Extrusion Temperature (ET), Feed Rate (FR), Layer Height (LH), Fan Power (F), and Bed Temperature (BT).1834A generic fractional factorial design is displayed in Table 1, where the sign (-) represents the low level and the sign (+) represents the high level of each printing parameter. The samples made for this research followed the runs presented in Table 1.

The samples made for this research followed the runs presented in Table 1. Each printing run was done in the same place of the printing bed with 3 samples printed at a time, repeating this process 3 times, resulting in a total of 144 samples for each printer.Table 1. Fractional factorial design for FFF process parameters effect on the extruded filament widthRun ET FR LH F BT + - - + + + + + + - - - - + + - + + + + - - + - - - + - + - + - - + + - - + + - - - + + - - + -1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16+ - + + + + - + - + - - - - + -+ - + - - - + + - + - + + + - -The values of the factorial design vary depending on the material and printer used. For ET and BT, the values depend on the material, these are presented in Table 2. For FR, LH, and F, the values depend on the 3D printer and nozzle to be used in the manufacturing process, these are presented in Table 3.

An AXIOCam ICc5 optical microscope, equipped with AxioVision SE64, was used to obtain the image and measurements of W. The top layer of the samples was the one analyzed to obtain the W measurements because it is not flat like the first layer that adheres to the printer bed.To allow a single filament to be observed under the microscope, it was used a magnification of 50x to analyze the W. 3 measurements were made along the length of the filament, as seen in Figure 2A, to measure the variation of W in the whole infill structure. The units of the measurements given by the microscope software were micrometers [𝜇𝑚], as represented in Figure 2B. The results of the OM were analyzed using MINITAB® as a statistical software program to obtain the regression models and residual analysis presented next.Figure 2.

The results of the OM were analyzed using MINITAB® as a statistical software program to obtain the regression models and residual analysis presented next.Figure 2. Image of a single filament subjected to OM.Results and DiscussionThe W results from optical microscopy were plotted using the means of the samples measured in each run configuration, these results are presented in Figure 3. Two regression analyses were done to obtain the parameter effects and interactions on W. The first regression was done to detect the parameters and the interactions that have a significant effect on the response. The second regression was done using only the parameters and interactions with significant effects obtained in the first regression. This results in a reduced number of parameters used in the models but maintaining a 𝑅2 higher than 70%, meaning that at least 70% of the process can be described with the parameters used.1836Figure 3.

Filament width variation OM results of samples made in Delta printer and Cartesian printers with A. PLA, B. ABS, and C. PC.PLA resultsSamples of PLA made in a Cartesian printer showed a range in W of 131.96𝜇𝑚 and an overall average in W of 359.13𝜇𝑚 as seen in Figure 3A as blue dots. The model, Equation (1), for PLA in Cartesian printer considers LH, LH-ET, LH-FR,FR-BT because they have the greatest impact on filament width response, obtaining a 𝑅2 = 96.89%, and the residuals accomplished normality, homoscedasticity and independency, as shown in Figure 4A.𝑊𝑃𝐿𝐴𝐶𝑎𝑟𝑡𝑒𝑠𝑖𝑎𝑛 = 359.13 − 53.93 𝐿𝐻 + 3.29 𝐸𝑇 ∗ 𝐿𝐻 − 3.25 𝐹𝑅 ∗ 𝐿𝐻 + 3.09 𝐹𝑅 ∗ 𝐵𝑇Samples of PLA made in Delta printer showed a range in W of 96.73 𝜇𝑚 and an overall average in W of 519.33 𝜇𝑚, as seen in Figure 3A as black diamonds.

The model for PLA in Delta printer, Equation (2), considers LH, FR, and LH-FR because they have the greatest impact on filament width response, obtaining a 𝑅2 = 94.61%, and residuals accomplished normality, homoscedasticity and independency, as shown in Figure 4B.𝑊𝑃𝐿𝐴𝐷𝑒𝑙𝑡𝑎 = 519.33 − 38.39𝐿𝐻 − 4.46 𝐿𝐻 ∗ 𝐹𝑅1837(1)(2)Figure 4. Residual analyses of W results of PLA samples made in A. Delta printer and B. Cartesian printer.ABS resultsSamples of ABS made in a Cartesian printer showed a range in W of 475𝜇𝑚 and an overall average in W of 659.7𝜇𝑚 as seen in Figure 3B as blue dots.

Delta printer and B. Cartesian printer.ABS resultsSamples of ABS made in a Cartesian printer showed a range in W of 475𝜇𝑚 and an overall average in W of 659.7𝜇𝑚 as seen in Figure 3B as blue dots. The model for ABS in Cartesian printer, equation (3), considers LH, BT, ET-F, and FR-BT because they have the greatest impact on filament width response, obtaining a 𝑅2 = 95.54%, and the residuals accomplished normality, homoscedasticity, and independency, as shown in Figure 5A.𝑊𝐴𝐵𝑆𝐶𝑎𝑟𝑡𝑒𝑠𝑖𝑎𝑛 = 659.66 − 195.45 𝐿𝐻 − 17.12 𝐵𝑇 + 18.41 𝐸𝑇 ∗ 𝐹 − 16.55 𝐹𝑅 ∗ 𝐵𝑇Samples of ABS made in Delta printer showed a range of 124.14𝜇𝑚 and an overall average of 520.65𝜇𝑚 as seen in Figure 3B as black diamonds.

The model for ABS in Delta printer, Equation (4), considers LH, F, BT and F-BT because they have the greatest impact on filament1838(1)response, obtaining a 𝑅2 = 81.77%, andwidth homoscedasticity, and independency, as shown in Figure 5B.residuals accomplished normality,𝑊𝐴𝐵𝑆𝐷𝑒𝑙𝑡𝑎 = 672.6 − 893.9 𝐿𝐻 + 69.4 𝐹 + 1.161 𝐵𝑇 − 1.223𝐹 ∗ 𝐵𝑇(2)Figure 5 Residual analyses of W results of ABS samples made in A. Delta printer and B. Cartesian printer.PC resultsSamples of PC made in Cartesian printer showed a range of 606.11 𝜇𝑚 and an overall average of 684.5 𝜇𝑚 as seen in Figure 3C as blue dots. When the factorial analysis was done, the residual analysis states that the residuals do not have normal behavior. An outlier test was performed, and a single outlier was deleted from the data, then the residual analysis of the factorial1839regression accomplishes normality, homoscedasticity, and independence.

The model for PC in the Cartesian printer, equation (5), considers LH, FR, FR-F, FR-LH and FR-BT because they have the greatest impact on filament width response, obtaining a 𝑅2 = 84.2%, and residuals accomplished normality, homoscedasticity and independency, as shown in Figure 6E.𝑊𝑃𝐶𝐶𝑎𝑟𝑡𝑒𝑠𝑖𝑎𝑛 = 684.5 − 40.9 𝐹𝑅 − 191.5 𝐿𝐻 + 39 𝐹𝑅 ∗ 𝐿𝐻 + 45.2 𝐹𝑅 ∗ 𝐹 − 38 𝐹𝑅 ∗ 𝐵𝑇Figure 6 Residual analyses of W results of PC samples made in A. Delta printer and B. Cartesian printer.Samples of PC made in Delta printer showed a range of 155.88 𝜇𝑚 and an overall average of 543.6 𝜇𝑚 as seen in Figure 3C as black diamonds.

Delta printer and B. Cartesian printer.Samples of PC made in Delta printer showed a range of 155.88 𝜇𝑚 and an overall average of 543.6 𝜇𝑚 as seen in Figure 3C as black diamonds. The model for PC in Delta printer equation (6), considers LH, BT, FR-F, ET-F, and F-BT because they have the greatest impact on filament width residuals accomplished normality, homoscedasticity and independency, as shown in Figure 6F.response, obtaining a 𝑅2 = 72.74%, and1840(3)𝑊𝑃𝐶𝐷𝑒𝑙𝑡𝑎 = 546.17 − 36.7 𝐿𝐻 + 14.85 𝐵𝑇 − 26.62 𝐸𝑇 ∗ 𝐹 + 28.81 𝐹𝑅 ∗ 𝐹 − 22.24 𝐹 ∗ 𝐵𝑇Response optimizationWhen regression models were obtained, they were summited to a response optimization. The optimization was specified to find the value of W closest to the diameter of the extruder used in each printer, 400 𝜇𝑚 for the Cartesian printer and 500 𝜇𝑚 for the Delta printer. Table 4 shows the results of the different printing parameters that were fitted on the response optimization process.Table 4.

Table 4 shows the results of the different printing parameters that were fitted on the response optimization process.Table 4. Printing parameter values approximating the predicted W to the diameter of the extruder used in each printer.Printer MaterialDeltaCartesianPLA ABS PC PLA ABS PCET [°𝐶] - - 250 215 230 -FR [𝑚𝑚3/𝑠] 720 - 720 1800 1800 1800LH [𝑚𝑚] 0.3 0.3 0.3 0.1 0.2 0.2F0 1 - 1 0BT [°𝐶] - 80 80 60 80 80W prediction [𝜇𝑚] 556.89 517.7 501.88 416.17 421.17 407.8In Delta printer, LH is the only parameter that is included in the regression models for PLA, ABS and PC. In Cartesian printers LH, FR and BT are included either as individual effect or as parameter interactions. All the regression analysis showed that LH has the largest impact on W, the lower LH, the higher W, as showed in Figure 7. As in Nuñez et al.[54], the higher LH selected gave the most accurate dimensional behavior.Figure 7. LH effect on extruded filament width. A.

As in Nuñez et al.[54], the higher LH selected gave the most accurate dimensional behavior.Figure 7. LH effect on extruded filament width. A. High LH decreases the W and B. Low LH increases the W.When the printers are compared, the Cartesian printer results showed higher standard deviations in W of ABS and PC, 206.7 and 229.4, respectively, in comparison with the Delta printer standard deviations, 49.10 and 69.8, for ABS and PC, respectively. Meaning that the dispersion of extruded filament width of ABS and PC is closer to the nozzle diameter in Delta printer than in Cartesian printer.1841(4)The Cartesian printer regression models have in common LH, BT, and FR as the printing parameters that most affect W.

The regression models obtained for the Delta printer have in common LH as the printing parameter that affects W in all the thermoplastics used.ConclusionsA fractional factorial design of experiments was used to find the effects of the parameters ET, LH, BT, F, and FR on filament width (W). It also helped to find predictive models for extruded filament width, using the least number of parameters and interactions between them, as well as the values of these parameters to have filaments with a thickness similar to the size of the extruder of the printers and thus obtain a homogeneous structure that has the same properties throughout the piece. Not having irregularities or variations will help to have a structure without stress concentrators that affect the result of any mechanical property at the time of performing mechanical tests on products made by FFF.The effect of layer height needs to be further analyzed to improve the control of FFF product dimensions.

Laser Based Powder Bed Melting Process offers design flexibility while enabling weight reduction due to topological optimization by substitute conventional design and manufacturing routes. Especially aluminum alloys are highly sensitive to oxygen and hydrogen impurity during the process. Due to this, it is mandatory to control precisely and hold a low oxygen level while processing. Hereby it is possible to avoid any negative impact on the final mechanical properties as tensile strength and fatigue resistance. For the investigations, aluminum powder (Al-Si-Mg) is used. This paper will present the influence of the oxygen level during processing on the final mechanical properties of the part.

For the investigations, aluminum powder (Al-Si-Mg) is used. This paper will present the influence of the oxygen level during processing on the final mechanical properties of the part. Even if pure Argon is commonly used to inert the chamber, different sources of oxygen like powder, equipment and gas supply have a negative impact and will be investigated.[1]–[4]IntroductionAdditive Layer Manufacturing (ALM) offers compared to conventional design and manufacturing routes benefits in cost reduction, design flexibility and the possibility of lead time reduction. An ALM related process is the Laser Based Powder Bed Melting Process (LBM). Aeronautic materials like titanium and aluminum alloys are already established for the LBM process. The serial production of LBM parts for titanium alloys is planned on particular parts within the next years.

The serial production of LBM parts for titanium alloys is planned on particular parts within the next years. The uses of aluminum alloys for the LBM process show a high sensitivity and complexity for welding, an affinity to hydrogen porosity and oxidation. Due to this further investigations are substantial.[4], [5] The build process and quality is mainly influenced by parameters like scanning speed, hatch spacing, laser power and layer thickness. Also the powder quality and process gas are directly influencing the mechanical properties of the produced part. Due to a good weldability and lightweight, aluminum silicon magnesium alloys are commonly used for aeronautic applications. [1], [3], [6]–[8] The aim of this paper is to investigate the influence of the powder quality and process gas. Current situation show that the majority of the ALM machines run with a higher value than 1000 ppm process gas contamination by oxygen and moisture.

Current situation show that the majority of the ALM machines run with a higher value than 1000 ppm process gas contamination by oxygen and moisture. For a precisely measurement of oxygen contamination and controlled process atmosphere the existing LBM machine is improved by an external cabinet developed by Linde AG and Airbus Group Innovations. This enables a reliable oxygen measurement as well as active control of the oxygen contamination in the process chamber. Two powders are processed under three different process gas levels for a defined parameter set: > 1000 ppm, 500 ppm and 30 ppm. The chemistry is measured for the raw material (powder) and molten material. To examine the mechanical properties tensile testing are accomplished.419ExperimentalI.Powder characterizationFor the investigations two AlSi10Mg powders supplied by two different powder suppliers are examined. In the following the powders are referenced as powder A and B. The atomization method is unknown.

In the following the powders are referenced as powder A and B. The atomization method is unknown. Powder delivery requirement for the particle size distribution is between 20 to 63 µm. Further information are investigated in former investigations of D. M. Bauer et al [9]. A scanning electron microscope (Zeiss Auriga FIB-SEM) is used to describe the powder morphology in terms of skin deep shape, satellites appearance, agglomeration rate and surface texture. Shown in Figure 1.Figure 1: SEM pictures of powder A (left) and powder B (right)Particle size distribution (PSD) is done with an optical measurement system (Malvern Morphologi G3), which analyzes the particles by a camera. A defined powder volume is dispersed by pressurized air onto a glass plate. The particles are characterized by the commonly used values D10, D50 and D90. Also a sphericity (spherical, non-spherical) measurement is done. The value of rest is not possible to determine.

The particles are characterized by the commonly used values D10, D50 and D90. Also a sphericity (spherical, non-spherical) measurement is done. The value of rest is not possible to determine. Results of PSD are shown in Table 1.Table 1: Particle size distribution of powder A and B measured for powder income characterization.PowderA BD10 [µm] 25 32D50 [µm] 45 53D90 [µm] 65 79Spherical [%] 40 10Non Spherical [%] 59 87To examine the powder flowability a dynamically measurement system (Revolution Powder Analyser, Mercury Scientific Inc.) is used. A volume of 100 ml powder is filled into a drum between two glass fronts whereon a camera system detects the avalanches due to the rotation of the drum. For each powder five test series of 150 avalanches are measured to calculate an average angle.

For each powder five test series of 150 avalanches are measured to calculate an average angle. The results of powder flowability measurement are shown in Table 2.Table 2: Flowability measurement of powder A and B.PowderAverage angleGradientA B50 490.053 0.064420Rest [%] 1 3The chemical composition is measured by Inductively Coupled Plasma Optical Emission Spectroscopy (ICP - OES). The Silicon content is determined by gravimetric measurement. The oxygen content (amount of process related powder and molten oxidation) is measured by carrier gas hot extraction. The chemical composition of powder A and B are shown in Table 3.Table 3: Che mical composition of powder A and B.PowderA BAl [wt%] Rest RestSi [wt%] 9.78 9.88Mg [wt%] 0.31 0.31Mn [wt%] 0.001 0.002O [wt%] 0.027 0.095II.Gas characterizationGas supplyArgon is supplied for this investigation with 5.0 quality and is stored in liquid form outside of the building.

Welding gases are classifies and designates shielding gases in accordance to DIN EN ISO 14175:2008-06 (Welding consumables – Gases and gas mixtures for fusion welding and allied processes). Rest impurities are nitrogen and carbon dioxide.Table 4: Impurities in technical gases in accordance to DIN EN ISO 14175:2008 -6Gas quality4.6 4.8 5.0Argon [%] > 99.996 > 99.998 > 99.9990Oxygen [ppm] < 4 < 3 < 2H2O [ppm] < 5 < 5 < 3Gas analysis with Linde Gas ManagementLinde Gas and Airbus Group Innovations implemented a gas management system to monitor and regulate the oxygen level in the process chamber of the LBM machine. Therefore a nozzle is installed in the process gas flow. Therewith it is possible to analyze the process gas for oxygen and moisture content. Afterwards it is possible to regulate the oxygen level to a set point.

Therewith it is possible to analyze the process gas for oxygen and moisture content. Afterwards it is possible to regulate the oxygen level to a set point. Figure 2 shows the schematic structure of the implemented Linde Gas Management System.Figure 2: Schematic structure of the Linde Gas Management System. (Dp* = dew point)421Results and discussionCross sectionsFigure 3 shows the cross section for the standard oxygen controlled atmosphere for powder A. Close under the surface spherically shaped gas pores can be detected. Compared to the standard controlled process, Figure 4 shows the cross section of the 30 ppm controlled oxygen atmosphere. For the higher magnification the lack of fusion is clear to see. No gas pores close to the surface can be investigated.

For each condition 5 specimens are tested. Build direction is in horizontal direction (XY). The specimens are machined to the final tensile test geometry B4x20 in accordance to DIN 50125. Static tensile tests are performed at ambient condition on a universal testing machine (Zwick 1474 100kN). The amount of Si and Mg are directly influencing achievable maximum strength. As the results show the amount of precipitation hardening possible by the Mg partition which leads to the evolution of coherent Mg2Si particles by the above mentioned post built heat treatment will be an additional strength factor. On to powder material oxide contamination, it seems the oxygen contamination of the build chamber penalizes final material ductility and its plasticity.

Build direction is z-direction. Figure 7 shows the surface fracture with a fine microstructure. Pores can be seen clear, highlighted by the white circles, as detected for the cross sections, shown in Figure 3. Compared to the standard process the specimens build up in the 30 ppm oxygen controlled atmosphere show less pores as shown in Figure 8. The white highlighted areas show lack of fusion, detected and shown in Figure 4. For both conditions the elongation of around 7 % is confirmed by the honeycomb structure of the fracture.

The white highlighted areas show lack of fusion, detected and shown in Figure 4. For both conditions the elongation of around 7 % is confirmed by the honeycomb structure of the fracture. Powder B shows similar results with the difference of a higher porosity for the specimens build up in the 30 ppm oxygen controlled atmosphere.423Figure 7: Fracture analysis for te nsile specimen build in standard oxygen controlled atmosphere for powder AFigure 8: Fracture analysis for te nsile specimen build in 30 ppm oxygen controlled atmosphere for powder A424ConclusionFor the investigations AlSi(Mg) alloy powders of two different suppliers with different oxygen contents (270ppm and 950ppm) are used. The samples were processed under three different atmospheres (30ppm, 500ppm and uncontrolled[>1000ppm] of oxygen contamination). The results of the investigation show:Reduction of gas porosity for a controlled process gas atmosphere of 30 ppm.

The results of the investigation show:Reduction of gas porosity for a controlled process gas atmosphere of 30 ppm. However, consequently,an adjustment of process parameters is mandatoryOxide contamination of the build chamber appears to have a direct impact on the chemical compositionand thus on the mechanical properties of the final partOxide contamination seems to penalize final material ductility and its plasticityTo process AlSi(Mg) alloys in a reproducible manner, an accurate and precise oxygen monitoring and especially a controlling system are mandatory.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEFFECT OF PROCESS PARAMETER VARIATION ON MICROSTRUCTURE AND MECHANICAL PROPERTIES OF ADDITIVELY MANUFACTURED TI-6AL-4VJonathan Pegues1, Kelvin Leung2, Azadeh Keshtgar2, Luca Airoldi2, Nicole Apetre2;Nagaraja Iyyer2, Nima Shamsaei1,*1Laboratory for Fatigue & Additive Manufacturing Excellence, Department of Mechanical Engineering, Auburn University, Auburn, AL 368492 Technical Data Analysis, Inc., 3190 Fairview Park Dr., Suite 650, Falls Church, VA 22042Corresponding author: shamsaei@auburn.eduAbstractAs additively manufactured (AM) parts become viable options for various structural applications, it is essential to fully understand how various process parameters affect their subsequent mechanical behavior.

In this study, a process simulation package is utilized to model the heat transfer during the laser-based powder bed fusion (L-PBF) process. Along with mechanical properties, the model is capable of estimating porosity, density, and microstructural evolution during the rapid cooling phases associated with L-PBF. Simulation results were validated using experimentally measured mechanical properties of L-PBF Ti-6Al-4V specimens. Tensile tests and microscopy were conducted for determining the effects of process parameters (i.e. laser power, laser scan speed, hatch distance) on microstructure, strength, and ductility of fabricated parts. Predictions from process simulations were compared with experimental results.

laser power, laser scan speed, hatch distance) on microstructure, strength, and ductility of fabricated parts. Predictions from process simulations were compared with experimental results. Validated process simulation packages, used and developed herein, can provide a low-cost means for the quality control of AM parts.Keywords: Mechanical properties; Microstructure; Additive manufacturing; ICME; Titanium alloysIntroductionRecently additive manufacturing (AM) has gained much interest to be implemented in many structural applications across a wide variety of industries including biomedical and aerospace applications. It has been well established, however, that parts produced via AM are not comparable to their wrought counterparts showing varying strength and ductility with the later showing large reductions in the AM parts [1–4].

The differences in strength and ductility of AM parts when compared to their wrought counterparts is mostly attributed to the unique microstructural characteristics of the AM parts including grain structure, chemical composition, residual stresses, presence of micro-voids, lack of fusion, etc. On the other hand, tailoring purpose61specific materials through manipulation of process parameters to achieve a more desirable microstructure and mechanical properties is an intriguing aspect of the AM method [5]. Research has shown that many of these microstructural characteristics can be manipulated by adjusting the AM process parameters [6].Experimentally determining the optimal process parameters for a desired microstructural or mechanical property requires a significant time investment coupled with high experimentation costs. In order to reduce these time and experimental costs a Process Parameter Optimizer (PPO) tool was developed to optimize the more critical process parameters.

In order to reduce these time and experimental costs a Process Parameter Optimizer (PPO) tool was developed to optimize the more critical process parameters. The framework utilizes a sensitivity analysis method to determine the process parameters which are most capable of altering the microstructure and resulting mechanical properties. The proposed PPO tool consists of an additive manufacturing (AM) simulation module, a sensitivity analysis of process parameters, and an AM process optimization module which are accompanied by supporting experiments to validate the simulation model. The PPO tool is ultimately designed to receive desired microstructural or mechanical properties by providing on optimized set of parameters through simulation.Multi-scale AM Simulation ModuleA multi-layer multi-track L-PBF process of Ti-6Al-4V was simulated via a finite element (FE) model for the 3D heat transfer analysis and is shown in Figure 1.

The heating of the powder layer is modeled using a time dependent heat transfer process which is similar to a diffusionprocess and given in Equation 1 where Q, T, Cp, ρ, and k are the body heat source, the temperature, the specific heat, the density and the thermal conductivity of the material, respectively.𝜌𝐶𝑝𝑑𝑇 𝑑𝑡= 𝑘∇2𝑇 + 𝑄(𝑥, 𝑡)(1)A representative section (3 mm × 5 mm × 0.5 mm) of the AM part was simulated with each deposited layer having the dimensions of 1 mm × 5 mm × 0.1 mm. The convergence of the maximum temperature was predicted through two models with different mesh refinement at the center of the top surface. To achieve convergence, the average element size was set to approximately the laser beam radius. A Goldak double ellipsoid volumetric heat source profile [7] was used to model the laser input to the sample. Additionally, a subroutine developed by Technical Data Analysis, Inc.

A Goldak double ellipsoid volumetric heat source profile [7] was used to model the laser input to the sample. Additionally, a subroutine developed by Technical Data Analysis, Inc. (TDA) was implemented to model the motion of the laser scan path.62Figure 1: Simulation results for lack of fusion zones for scan velocities of a) 0.5 m/s and b) 0.6 m/s.One of the most influential microstructural characteristics for AM parts is the lack of fusion between subsequent layers. These void-like areas have very little volume and therefore do not result in a significant loss of density, however, their flat geometry results in greater stress concentrations than typical spherical voids. Model simulations were conducted with the specific intent of modeling the lack of fusion zones within an AM part.

Model simulations were conducted with the specific intent of modeling the lack of fusion zones within an AM part. Six AM process input parameters were selected to study their effects on final temperature and porosity, and include laser beam power (P), efficiency of the laser beam (η), laser beam radius (c), laser scan speed (v), ambient temperature (Tamb), and absorptivity of the material (α). To illustrate the heat transfer simulation, the temperature history of two selected controlled points in the simulation domain for v = 5 mm/s at different locations of the domain are presented in Figure 1. Figure 1 also shows the maximum temperature achieved during the simulation at two cross section locations. The lack of fusion (LoF) regions can be determined by the maximum temperature plots. Incomplete melting occurs when the maximum temperature achieved during the simulation is less than the melting temperature of Ti-6Al-4V.

Incomplete melting occurs when the maximum temperature achieved during the simulation is less than the melting temperature of Ti-6Al-4V. For convenience, the LoF regions are depicted in the figures by thicker contour lines. By sampling sections, the porosity of the AM part can be estimated by summing up the volume of the LoF regions. Results from the simulation show an increase in lack of fusion areas with increasing scan speed as shown in Figure 1.63Sensitivity AnalysisThe AM process can be affected by a multitude of build parameters; however, the effect of each parameter may be insignificant when compared to others. To reduce computational and experimental time costs, a sensitivity analysis was conducted to determine which parameters will show the greatest effects on the AM parts.TDA has developed a novel global sensitivity analysis method based on Sobol’s index and Latin Hypercube Sampling (LHS).

The proposed global sensitivity analysis approach varies all variables simultaneously through its entire range by LHS while considering the interaction between different parameters (such as laser power and beam width). The LHS method has the ability to recreate the input distribution through sampling in fewer iterations when compared with the Monte Carlo method. Fewer iterations for LHS is achieved by stratification of the input probability distributions. Sobol’s method uses the decomposition of variance to calculate the Sobol’s sensitivity indexes [8].

Both primary and complementary parameter spaces are needed to implement Equation 5.𝐷𝑖 = 𝐷 −1 2[𝑓(𝑥) − 𝑓(𝑥𝑖, 𝑥′−𝑖)]2 𝑑𝑥 𝑑𝑥 ′−𝑖 ≈ 𝐷 −1 2𝑁𝑁 ∑[𝑓(𝑥𝑘) − 𝑓(𝑥𝑖𝑘, 𝑥′ 𝑘=1−𝑖𝑘)]2(5)The total sensitivity effects are computed for all the chosen parameters using Equation 6.64𝑡𝑜𝑡 =𝐷𝑖1 2∫[𝑓(𝑥) − 𝑓(𝑥𝑖, 𝑥′−𝑖)]2 𝑑𝑥 𝑑𝑥 ′−𝑖 ≈1 2𝑁𝑁 ∑[𝑓(𝑥𝑘) − 𝑓(𝑥′ 𝑘=1𝑖𝑘, 𝑥−𝑖𝑘)]2(6)The parameters are then sorted according to their sensitivities.A Matlab code developed by TDA was used to implement the described approach to Sobol’s index. Three process parameters were selected as a focus on the global sensitivity analysis. These parameters include laser power, laser scan velocity, and laser beam width. A Monte Carlo Simulation was implemented for the simulation model with the process parameter ranges given in Table 1.

These parameters include laser power, laser scan velocity, and laser beam width. A Monte Carlo Simulation was implemented for the simulation model with the process parameter ranges given in Table 1. A total of 100 iterations were conducted for the select ranges with the maximum temperature being estimated via the heat transfer module previously mention and recorded for each iteration. Figure 2 shows the calculated results of the first order effects as well as the total effects for each parameter. From these results, it can be seen that the laser scan velocity was the most sensitive for the ranges given while the laser beam width showed the least sensitivity. It should be noted, however, that there is an intrinsic relationship between the laser power and laser beam width, however, compared to laser power the effect of scan velocity is much higher (Fig. 2).

2). It should be mentioned here that these specific results are limited to the TDA simulation data and the specified process parameters listed in Table 1. These results are not necessarily applicable directly to other case studies. It is recommended to implement a complete proposed global sensitivity analysis for each case study individually.Table 1: Process parameter values used for sensitivity analysis.L-PBF ParameterLaser Power (W) Laser Scan Velocity (m/s) Laser Beam Spot Size (m)Mean Value Lower Bound Upper Bound 252 1.08 4.5e-5280 1.20 5.0e-5308 1.32 5.5e-5Figure 2: First order and total effects for the given process parameters.65Experimental ResultsSpecimen Design and FabricationIn order to validate the model simulations, a series of tensile tests were conducted on a set of AM fabricated samples with varying process parameters.

A total of three process parameters were chosen to experimentally determine the effect each had on the mechanical properties of the AM parts. The process parameters chosen for this research was laser power, scan velocity, and hatch spacing. The process parameters can be separated into three categories in which two of the three parameters are held constant to produce a set of varying laser power, scan velocities, and hatch spacing as shown in Table 2.

The general set of parameters for Ti-6Al-4V provided with the AM machine is given as parameter set two, which is bolded in Table 2 and included as a base for each of the three categories.Table 2: Parameter sets used to produce the AM parts separated into their respective categories.ParameterLaser Power (W)Scan Speed (mm/s)Hatch Distance (m)125212001402280120014033081200140103361200140113641200140142808401401328096014042801080140228012001405280132014072801200126228012001406280120015416280120016866Specimens were fabricated using a powder bed based EOS M290 with a maximum laser power of 400W. The spherical Ti-6Al-4V alloy powder was supplied with a powder size greater than 15 m and less than 45 m. Specimens were built horizontally onto a titanium substrate as shown in Figure 3 with the given dimensions.

Specimens were built horizontally onto a titanium substrate as shown in Figure 3 with the given dimensions. Three to four samples were produced for each parameter set given in Table 2. The specimens were stress relieved before removal from the plate to remove any residual stress as result of the increased cooling rates associated with the L-PBF process. The stress relief was conducted at 700° C for 1 hr under an Ar environment and allowed to air cool. Upon removal, each specimen surface was lightly ground by hand in order to make more accurate cross-sectional measurements of the cross-section.Figure 3: Sample orientation on build plate and specimen dimensions.Tensile tests were conducted on a servo-hydraulic MTS testing frame with a load capacity of 100 kN. The tests were conducted in strain control with a strain rate of 0.001 s-1 until sufficiently past the yield stress (y) after which the test was continued in displacement control until final fracture.

Fracture surfaces were removed along with a portion of the gage section for further67microstructural analysis. The fracture surfaces were observed using a scanning electron microscope with an excitation voltage of 20 kV. The as-built and removed gage sections were prepared using standard metallographic procedures. The microstructure was reveled using a mixture of 92 ml of distilled water, 6 ml of nitric acid, and 2 ml of hydrofluoric acid with each sample being immersed for 20 – 30 s. The microstructure consisted of mostly alpha () and martensite (’) grains within elongated prior beta () grains. The ’ grains are readily observable by the acicular grains within the prior  grains as shown in Figure 4. The prior  grains show equiaxed grains parallel to the build direction (Fig. 4(a)) while showing more elongated grains perpendicular to the build direction due to epitaxial grain growth (Fig.

The prior  grains show equiaxed grains parallel to the build direction (Fig. 4(a)) while showing more elongated grains perpendicular to the build direction due to epitaxial grain growth (Fig. 4(b)).Figure 4: Microstructure of AM parts for a) longitudinal direction and b) cross section direction.Tensile BehaviorThe tensile properties were mostly similar across all process parameters with the only exception being elongation to failure and fracture strain as shown in Table 3.

AM parts produced by L-PBF typically show increases in yield stress (y) ultimate tensile strength (UTS) with decreasing ductility when compared to their wrought counterparts which was observed for the present investigation and shown in Table 3.68Table 3: Tensile properties for each parameter set.ParameterEyUTS Elong.ff#GPa MPa MPa%MPa %195.8108411333.4011703.282102.0107911302.8311683.313102.4111111693.8812133.78499.0106311123.8111533.65598.0107111274.0411583.65697.6107411364.6011834.35799.5106611335.3511864.981095.8105011233.0411562.951199.4106811374.2411834.0713100.3105211184.2711654.081497.3107311365.6511895.351694.9107311444.7211854.23Wrought (average)113.088095014.00//Computed Tomography ScanX-ray tomography was conducted on a Shimadzu Safire Medical Tomosynthesis system.

Three scans of middle regions of 3 coupons were performed and are shown in Figure 5. The tomography results show that, the most consistent production defects of these specimens are: (a) surface roughness, (b) voids, and (c) inclusions. The tomography results showed a much greater number of voids in sample #5 as shown in Figure 5. Along with this significantly greater number of voids present within the specimens compared to the other tested samples, was a much rougher surface, 33.42 μm compared to 24.55 (Sample # 3) and 16.94 (Sample #6). Interestingly, sample69set # 5 had a greater scan speed (1.32 m/s) compared to sample #3 and #6 (1.20 m/s).

Interestingly, sample69set # 5 had a greater scan speed (1.32 m/s) compared to sample #3 and #6 (1.20 m/s). As previously mentioned, the scan velocity showed the higher sensitivity and the simulation showed that increasing scan velocity would result in higher lack of fusion (void) area.Figure 5: Tomography results for three selected specimens.70FractographyThe fracture surfaces of selected samples were observed with a scanning electron microscope (SEM). Two characteristic surfaces were readily observable for all fractured specimens, a fibrous area and a ductile overload area. The fibrous area is shown in Figure 6 outlined in red and is characterized by transgranular quasi-cleavage fracture. The ductile overload area, shown outside the fibrous area, shows shear failure characteristics. In the fibrous zone a combination of ductile fracture and cleavage fracture is observed.

The ductile overload area, shown outside the fibrous area, shows shear failure characteristics. In the fibrous zone a combination of ductile fracture and cleavage fracture is observed. Striations consistent with fracture along the ’ phase are observable in the fibrous section of the fracture surfaces as shown in Figure 6 (blue). The ductile overload area shows a more elongated dimpled structure with the characteristic shear angle approximately 45° from the fibrous section as shown in Figure 6 (red).Figure 6: Fracture surface of selected tensile sample showing the fibrous area outlined in red.DiscussionTi-6Al-4V with an ’ microstructure shows much less ductility than it’s  +  counterparts [9] and behaves in a more similar manner to cast Ti-6Al-4V. As such, the L-PBF produced Ti- 6Al-4V parts with high ’ phase fractions would be expected to show much lower ductility than their wrought counterparts.

As such, the L-PBF produced Ti- 6Al-4V parts with high ’ phase fractions would be expected to show much lower ductility than their wrought counterparts. However, the high ’ microstructure does not explain the variation in ductility between the parameter sets observed in this investigation. Additive manufactured parts can be more susceptible to quasi-brittle fracture due to the type of defects intrinsic to the L-PBF process such as the lack of fusion between layers and voids throughout the parts. These defects can also result in lower ductility for AM parts when compared to their wrought counterparts. These effects can be readily observed when comparing the fracture toughness of the three specimens used71for tomography.

These effects can be readily observed when comparing the fracture toughness of the three specimens used71for tomography. Comparing just the tomography samples, it can be seen that the sample with the highest number of voids (sample #5) had the lowest ductility while the sample with the lowest number of voids (sample #6) had the highest ductility.The sensitivity analysis discussed previously predicted that scan velocity would have the most effect on the maximum temperature during the building process and therefore could show the most effect on the mechanical properties. The experimental validation was able to demonstrate this effect as shown in Figure 7 where, compared to laser power and hatch distance, changes in scan velocity showed much greater effects on the ductility. Increasing scan velocity results in a decrease in true fracture strain resulting in less ductility.

Increasing scan velocity results in a decrease in true fracture strain resulting in less ductility. The decrease in ductility is most likely a result of a higher fraction of ’ phase due to more rapid heating and cooling along the beam path for higher scan velocities. Research has shown that wrought Ti-6Al-4V alloys with high ’ phase fractions show much less ductility compared to their  + counterparts [9]. As such, additively manufactured Ti-6Al-4V parts by L-PBF with intrinsically high cooling rates resulting in higher ’ should show similar behavior. Continuing this logic, L-PBF produced parts with lower scan velocities should show a lower fraction of the ’ phase with an improvement in ductility compared to higher scan velocities.Laser power showed some effect on ductility, however, not as consistently or to the same effect as scan velocity.

Generally, higher laser power resulted in greater ductility, however, it can be seen in Figure 7(b) that the second highest laser power resulted in the lowest ductility. Hatch distance showed relatively no effect on the ductility of the AM parts. The lack of effect of hatch size on ductility is not surprising since it does not affect the volumetric energy associated with the interaction of the beam and powder. There should be some slight effect on the cooling rate for a given point on the build, however, compared to the effect of scan velocity and laser power, hatch distance would not be expected to show large variations in temperature.ConclusionsIn summary, a multi-scale AM simulation module developed by TDA was implemented to simulate the effect of varying process parameters on the AM produced parts. A sensitivity analysis was then conducted to determine which of the process parameters in the AM process could affect the mechanical properties of the AM parts the most.

A sensitivity analysis was then conducted to determine which of the process parameters in the AM process could affect the mechanical properties of the AM parts the most. It was predicted that the AM parts would show more sensitivity to scan velocity than laser power. Simulation showed that increasing scan velocities would result in greater porosity and lack of fusion areas. Tomography analysis of the AM produced parts showed that the sample with the highest scan velocity contained a much higher number of voided areas than lower scan velocities. Further experimental analysis then showed that the ductility of the AM parts was more sensitive to scan velocity compared to laser power validating the sensitivity analysis.

Further experimental analysis then showed that the ductility of the AM parts was more sensitive to scan velocity compared to laser power validating the sensitivity analysis. Additionally, changes in the process parameters showed little effect on y and UTS while showing more effect on the ductility of the AM parts.72Figure 7: Process parameter effects on true fracture strain.73AcknowledgementsThe funding from the Naval Air Systems Command (NAVAIR) is greatly appreciated. The authors would like to also thank Professor Andrew Makeev for the tomography results which were carried out at the Advanced Materials and Structures Lab (AMSL) at University of Texas at Arlington.References[1] N. Shamsaei, A. Yadollahi, L. Bian, and S. M. Thompson, “An overview of Direct Laser Deposition for additive manufacturing; Part II: Mechanical behavior, process parameter optimization and control,” Addit. Manuf., vol. 8, pp. 12–35, Oct. 2015.[2] A. J. Sterling, B. Torries, N. Shamsaei, S. M.

fect of Process Parameters and Heat Treatment on the Microstructure and Mechanical Properties of SLM-built HY100 steelJ.J.S. Dilip*, Brent Stucker * and Thomas L. Starr**Department of Industrial Engineering, **Department of Chemical Engineering, J.B. Speed School of Engineering, University of Louisville, Louisville, KY 40292AbstractHY100 is a high strength low alloy steel used for naval and pressure vessel applications. In general, the alloy is used in the quenched and tempered condition. In the present work, fully dense metallic samples were produced from HY100 pre-alloyed powders using selective laser melting (SLM). Test samples were built with varying process parameters (scan speed and laser power). The SLM-built samples were given direct tempering treatment and a standard quench and temper heat treatment. Tensile properties of the samples were evaluated in direct temper, and quench and temper conditions.

Tensile properties of the samples were evaluated in direct temper, and quench and temper conditions. The study investigates the influence of process parameters and heat treatment on the microstructure and mechanical properties of SLM-built HY100 steel.IntroductionAdditive manufacturing involves layer-by-layer fabrication of a three-dimensional part based on computer-aided design data. It involves mathematically slicing a CAD model of the part to be produced into a number of horizontal layers and systematically creating these layers one over another, from bottom to top, to produce the part. There are a number of commercially available additive manufacturing processes such as Laser Engineered Net Shaping (LENS), Selective Laser Melting (SLM), and Electron Beam Melting (EBM), which can produce complex-shaped end-use metallic parts [1].

In all these processes material addition is achieved through melting and solidification.Selective Laser Melting (SLM) is a powder bed fusion-based additive manufacturing process in which successive layers of metal powder particles are melted and solidified on top of each other by a high intensity laser beam [2]. SLM of materials is generally accompanied with high cooling rates due to the short dwell time and high thermal gradients. These high cooling rates during SLM may result in the formation of finer microstructural features when compared to traditional manufacturing methods [2,3]. The process of layer-by-layer deposition causes each deposited layer to experience multiple thermal cycles during the build which in-turn can result in varied microstructures. Microstructures are governed by the processing parameters and in turn, control the part mechanical properties [3,4].In the present study, HY100 pre-alloyed powders were used to produce 3D metallic parts using SLM.

The alloy usually gains its required properties by heat treatment, which initially involve austenitizing at high temperature and quenching to form fully martensitic microstructures (hard and brittle). The high hardness of quenched martensite is mainly due to high carbon concentration in the lattice of martensite [5]. To obtain acceptable levels of ductility and toughness, the alloy is subjected to tempering. Isothermal holding at the tempering temperature causes the carbon in the martensite lattice to precipitate out as carbide particles within the martensite phase. The formation of these carbides results in significant reduction in hardness and improved ductility and toughness with an acceptable loss in yield strength [6]. The final microstructure after quench-and-temper condition consists of tempered martensite and/ or bainite phases [7].

The final microstructure after quench-and-temper condition consists of tempered martensite and/ or bainite phases [7]. The microstructure is tuned such that it gives a high yield strength (690 MPa, min.) while maintaining good toughness and ductility [8].Essentially no information is available on SLM of HY100 steel in the open literature. Microstructure evolution in this steel during SLM can be complicated because of the multiple thermal cycling effects. Hence, the parts in as-built condition may develop some microstructure-property problems. Therefore, it may be necessary to employ a suitable heat treatment for achieving satisfactory part mechanical properties.

Therefore, it may be necessary to employ a suitable heat treatment for achieving satisfactory part mechanical properties. The objectives of the present work are:(i) Optimization of the processing parameters to produce fully dense metallic samples from HY100pre-alloyed powders.(ii) Studies on mechanical properties of the HY100 SLM samples in various heat treatmentconditions.Experimental methodsGas atomized HY100 pre-alloy steel powder (Sandvik Osprey) were used in the present study. The nominal composition of the HY100 powders is presented in Table 1. The powder particle size and distribution were measured using a Microtrac S3000 particle size analyzer. Further, SEM observations for the size and morphology of the powders were also performed.Table 1.

Further, SEM observations for the size and morphology of the powders were also performed.Table 1. Chemical composition of HY100 steel powdersElement C 0.1- wt.% 0.22Ni 2.6- 3.5Cr Mo 0.27- 0.631.29- 1.86P 0.02 max.S 0.02 max.Mn 0.1- 0.45Si 0.12- 0.38V 0.03Ti 0.02Cu Fe 0.25 Bal.The processing of HY100 powder involves two steps: fabrication in the SLM machine and post-build thermal treatment. Samples were produced using an EOS M270 Direct Metal Laser Sintering (DMLS) machine in a nitrogen purging atmosphere. The protective gas allows maintaining the oxygen content to low levels to prevent the oxidation of HY100 powders during part fabrication. Initially, small cubes (10 mm X 10 mm X 10 mm) were fabricated using several parameter sets in order to study the effect of processing parameters on the density of SLM-processed parts and to select “optimal” parameters to fabricate tensile specimens.

Seven parameter sets were created by varying laser power, scan speed and scan offset distance. The combination of parameters was constrained to three energy densities. The parameters used are listed in table 2.Table 2. Process parameters for fabricating HY100 steel samples.Parameters #1 2 3 4 5 6 7Laser Power (W)195 195 175 195 195 195 195Hatch Distance (μm) 100 120 100 120 100 120 100Speed (mm/s)833 700 750 765 920 625 750Layer thickness (μm) 40 40 40 40 40 40 40Energy density (J/mm3) 58 58 58 53 53 65 65A mild steel plate with dimensions 300 mm × 300 mm × 25 mm was used as the build platform. Flat tensile specimens (ASTM E-8) were fabricated using the optimal parameter set. Fig.1 shows the base plate with the tensile samples and small cubes.

Flat tensile specimens (ASTM E-8) were fabricated using the optimal parameter set. Fig.1 shows the base plate with the tensile samples and small cubes. As can be seen from Fig.1, the samples were not precisely oriented along the x and y axes, so throughout the rest of this paper the samples that were most nearly aligned with the x and y axes are labeled XZ and YZ build orientations respectively. These builds were used for carrying out various heat treatments (Table 3) and tensile tests.Microstructural characterization was carried out on the as-built samples and on the heat treated samples. The samples were sectioned using a low speed saw. The cut samples were mounted in resin and polished using SiC sandpaper. Final polishing was carried out using 1 μm diamond suspension. The samples were etched using 4% Nital reagent. The cross-sectional microstructure of the unetched and etched conditions were examined by optical microscopy (OM) and SEM.

The samples were etched using 4% Nital reagent. The cross-sectional microstructure of the unetched and etched conditions were examined by optical microscopy (OM) and SEM. Rockwell hardness measurements were carried out on the as-built and heat treated samples. The reported values are the average of five indentations measured on the sample. Hardness results are presented in table 3.Tensile tests were carried out for specimens fabricated in XZ and YZ orientations with as-built and heat treated tensile samples. The tensile test results such as yield strength, ultimate tensile strength, and percentage strain-to-failure were reported. The percentage strain-to-failure was measured using an extensometer clipped to the gage section of the test specimen. Fracture surface examination of the tensile tested samples was also performed using SEM.Fig.

Fracture surface examination of the tensile tested samples was also performed using SEM.Fig. 1 Photograph showing a build plate with flat tensile samples (YZ) and small cubesResults and DiscussionFig.2 a & b show the morphology of the powder particles. The diameters of individual particles were measured from several different SEM micrographs to estimate the size and distribution of the particles. It was estimated that particles have a size distribution between 5 μm and 40 μm. The powders were spherical in shape, which is an important powder characteristic for good flowability. A good flowability is necessary for better powder layer deposition and also ensures that the powders can be dispensed properly [9]. SEM shows typical dendritic morphology features on the surface of the particles. The larger particles were observed to have satellite particles attached. These small particles form due tofragment micro droplets sticking to the larger particles.

The larger particles were observed to have satellite particles attached. These small particles form due tofragment micro droplets sticking to the larger particles. The mean particle diameter is ~30 μm with a standard deviation of 10 μm as measured by particle size analysis. The results from particle size analysis show a bi-modal distribution of the powder particles (Fig.2b).Fig. 2 (a) Scanning electron micrographs of the HY100 powders (b) Particle size distribution plotFor porosity measurement the cubes were metallographically prepared and examined in a polished condition using optical microscopy. Porosity in the sample was quantified using image analysis software. Fig.3 shows typical cross-section microstructures of SLM built cubes with different processing parameters. The dark irregular regions in the micrographs are pores (dark arrow) (Fig.3a).

Fig.3 shows typical cross-section microstructures of SLM built cubes with different processing parameters. The dark irregular regions in the micrographs are pores (dark arrow) (Fig.3a). For each value of energy density, the porosity was not significantly different for different scan speed, power and distance parameters – at least within the range of values used. The deposits produced using an energy density of less than 60 J/mm3 showed some porosity (1-2%), but those produced at higher energy densities were fully dense (Fig.3b). All SLM built HY 100 samples contained very fine non- metallic oxide particles in the matrix (fine dark spots in the micrograph Fig.3b). After these series of experiments were performed to establish suitable parameters for fully dense parts, processing parameter combination # 6 was used for further study.Fig.

After these series of experiments were performed to establish suitable parameters for fully dense parts, processing parameter combination # 6 was used for further study.Fig. 3 Typical cross-sectional micrograph of an as polished sample (a) with significant amount of porosity (parameter set #5) (b) fully dense (parameter set #6)Energy density is an important factor that affects the properties of as-built parts fabricated by SLM processing. The energy density E (J/mm3) in SLM is expressed in equation (1), where P is laser power (W), v is scan speed (mm/s), h is hatch spacing (mm) and t is layer thickness (mm) [10].E = P / v*h*tEnergy density directly indicates the energy input into to the process (ignoring any base plate or other bulk heating). A higher laser power provides more energy density for good melting and spreading of the melt pool.

A higher laser power provides more energy density for good melting and spreading of the melt pool. However, fusion defects or porosity can result if the scan speed is too high for a given laser power (due to insufficient laser-material interaction time). Similarly, too high a hatch spacing can cause fusion defects between individual scan lines or tracks. It should be noted that scan speed, hatch spacing and laser power are mutually dependent and are to be optimized considering their combined effects during SLM [11]. In the present study, the parameters with higher energy density appear to be near-optimum for producing fully dense samples from HY100.Fig. 4 Micrographs of an as-built fully dense HY100 steel sample (a) Optical and (b) SEM.An optical micrograph of a vertical cross-section of a fully dense SLM fabricated cube sample is presented in Fig.4a. The micrograph shows multiple layers clearly delineated by the dark etch bands along the fusion boundaries.

The micrograph shows multiple layers clearly delineated by the dark etch bands along the fusion boundaries. During SLM the powders are exposed to high intensity laser energy, which causes the powder temperature to rise above the melting point. As the laser passes away the molten pool of metal solidifies along the line of laser scan. The laser beam also causes partial melting of the below/adjacent layers, thereby allowing fusion welding of the layers. The alloy HY100 initially solidifies as delta ferrite directly from the liquid phase and subsequently transforms to austenite. From the continuous cooling transformation curve of the alloy, the high temperature austenite transforms to martensite as the cooling rates experienced in the process are steep [7]. The solidified track of alloy will now consist of untempered martensite. The reported Ms and Mf temperatures for the alloy are much higher than room temperature.

The solidified track of alloy will now consist of untempered martensite. The reported Ms and Mf temperatures for the alloy are much higher than room temperature. It is well-established that Ms temperature determines the microstructure and substructure formed during fast cooling. In the present alloy, lath martensite is expected to form due to higher Ms temperature [5]. Fig.4b (region indicated by arrow in Fig.4a) shows a high magnification SEM micrograph of the as deposited sample revealing lath morphology of martensite.Since the part fabrication takes place by layer-by-layer melting and solidification of the powders, the deposited layers underneath experience complex multiple thermal cycling effects. This scenario is equivalent to multi-pass welding of the material. Hence, it can be anticipated that during layer-wise(1)deposition, the regions which are reheated to high enough temperatures can experience martensite tempering [12].

Hence, it can be anticipated that during layer-wise(1)deposition, the regions which are reheated to high enough temperatures can experience martensite tempering [12]. However, in the present situation the thermal cycles are very brief, which can result in partial or slight tempering of martensite. In other words, it is quite unlikely that complete tempering of martensite would result during laser processing. Hence, the part in the as-built condition is unsuitable for direct application due to the hard untempered martensitic microstructure. Also, laser deposited parts are prone to residual stress accumulation during the process [13]. Thus, it is essential to determine proper heat treatments for achieving desired part mechanical properties.The as-built cubes were subjected to different heat treatments and Rockwell hardness tests (Table 3).

In one set of experiments the as-built cubes were given direct tempering and in another set, initial stress relief followed by full heat treatment (by austenize and quench, and then tempering). The stress relieving treatment was carried out at a maximum temperature of 600ºC for 1 hour. Table 3 summarizes the overall heat treatment effects on hardness.The MIL-S-16216 standard specification emphasizes an austenizing temperature of 900ºC for 1 hour followed by a quenching and tempering temperature not less than ~620ºC for HY100.The MIL standard leaves selection of tempering temperature to the manufacturer for meeting the desired mechanical properties [8]. It is well known from literature that tempering causes reduction in hardness of martensite due to the removal of carbon from the martensite lattice. The carbon reacts with Fe, Cr, and Mo and forms carbides within the martensite phase.

The carbon reacts with Fe, Cr, and Mo and forms carbides within the martensite phase. The removal of carbon and formation of carbide particles will generally result in improved toughness / ductility with an acceptable loss in strength and hardness. This quench and temper process creates a tempered martensitic structure. The hardness measurements show reduction in hardness due to tempering of martensite. As expected, the hardness of the heat treated samples was influenced by the tempering temperature. It was quite interesting to note that the hardness of the direct temper and full heat treatment samples did not show large differences. The hardness of the as-built sample was similar to the direct quench sample, which further confirms that during multilayer deposition tempering of martensite has no considerable effect.Table 3.

The hardness of the as-built sample was similar to the direct quench sample, which further confirms that during multilayer deposition tempering of martensite has no considerable effect.Table 3. List of heat treatments carried out on SLM HY100 samples and their corresponding hardness.Heat treatmentAs-built HY 100 900ºC-1h-WQ* Direct temper 620ºC-2h-AC** 900ºC-1h-WQ* + 620ºC-2h-AC** Direct temper 650ºC-2h-AC** 900ºC-1h-WQ* + 650ºC-2h-AC** Direct temper 670ºC-2h-AC** 900ºC-1h-WQ* + 670ºC-2h-AC**HRc (avg. of 5 readings) 37 38 26 21 22 22 20 18WQ : Water Quench, ** AC : Air CoolThe heat treated samples were examined under SEM and micrographs are presented in Fig.5. The micrographs show tempered martensite and bainite. It was very difficult to deduce quantitative differences in microstructures between the direct and fully heat treated samples. Nevertheless, high temperature tempered samples (670ºC) show coarser microstructural features.

Nevertheless, high temperature tempered samples (670ºC) show coarser microstructural features. Further detailed highermagnification studies by TEM are necessary in order to fully understand the nature of carbides, their size and distribution in the microstructure and their effects on mechanical properties.Fig. 5 SEM micrographs of HY100 steel in different heat treatment conditions (a) Direct temper at 620ºC (b) Quench-and-temper at 620ºC (c) Direct temper at 650ºC (d) Quench-and-temper at 650ºC (e) Direct temper at 670ºC (f) Quench-and-temper at 670ºC.It is well agreed and documented that hardness bears a direct correlation to the strength of material [14]. A reduction in hardness indicates softening of martensite and also the degree of tempering. Therefore, heat treatments on tensile samples were carried out using direct temper and full heattreatment. Three different tempering temperatures (620, 650 and 670ºC) were used.

Therefore, heat treatments on tensile samples were carried out using direct temper and full heattreatment. Three different tempering temperatures (620, 650 and 670ºC) were used. Tensile tests were carried out on the heat treated samples and results are given in Table 4.Table. 4 Tensile test results for HY100 steel deposits in different heat treatment conditions.Heat treatment0.2 % Y.S (MPa)UTS (MPa)Strain to failureHardness HRcAs-built (YZ)1172 1353637As-built (XZ)1178 13254.5--Direct temper 620ºC-2h-AC9159262.526900ºC-1h-WQ + 620ºC-2h-AC684756721Direct temper 650ºC-2h-AC7508804.522900 ºC-1h-WQ + 650ºC-2h-AC6586963.522Direct temper 670ºC-2h-AC704886720900ºC-1h-WQ + 670ºC-2h-AC586750718The tensile test results show promising strength values in the heat treated specimens. The as deposited samples with different orientations showed similar mechanical properties.

The as deposited samples with different orientations showed similar mechanical properties. This was expected since the scan pattern variation in these samples should normalize properties in the XY plane. It is expected, however, that samples in the Z orientation may have differing mechanical properties. The heat treated samples showed a reduction in strength levels as the tempering temperature was increased. It was interesting to note that direct temper samples gave good yield strength values. Therefore, it can be deduced that HY100 steel SLM fabricated parts can be subjected to direct tempering, thus eliminating the austenizing and quenching steps. The ductility of the as-built and heat treated samples was quite limited. SEM fractography studies of the fracture surfaces revealed (Fig.6) typical dimple rupture failure mode with transgranular ductile failure in the direct tempered and fully heat treated deposits.

SEM fractography studies of the fracture surfaces revealed (Fig.6) typical dimple rupture failure mode with transgranular ductile failure in the direct tempered and fully heat treated deposits. The sizes of the dimples were observed to be fine and shallow in the fracture surfaces. The size and shape of the dimples indicate the formability of the material, and finer dimples indicate lower elongation values [15]. The elongation values obtained are somewhat less than for the MIL-S-16216 specification. The phenomenon of lower elongation than expected was reported by several researchers for various additively manufactured alloys [16,17,18]. At present, the reason for lower ductility in the present HY100 alloy made using SLM is not fully clear. However, lower ductility can be partly attributed to the presence of fine nonmetallic oxides (Fig.7) and contribution from secondary hardening by precipitation of Molybdenum carbide during tempering [5].

These oxides were confirmed from EDS analysis as complex oxides of Al, Si and Mn. The presence of fine oxide particles will have significant effect on fracture initiation and propagation during tensile testing [19].Fig. 6 SEM fracture surface of a direct tempered sample (650ºC-2h-AC). The fracture features exhibit a typical dimpled rupture mode of failure.Fig. 7 SEM micrograph of the etched cross- section of the as-built, fully dense HY100 steel sample showing fine nonmetallic oxide particles.ConclusionsIn the current study, selective laser melting experiments were conducted using HY100 steel powders. Microstructural and mechanical properties of the HY100 SLM deposits were carried out. The following conclusions can be drawn from the present work:HY100 can processed using selective laser melting. Parameters for fully dense parts are possible.Thermal cycles involved in SLM have an effect on HY100 microstructural development, resulting in non-traditional microstructures.

Parameters for fully dense parts are possible.Thermal cycles involved in SLM have an effect on HY100 microstructural development, resulting in non-traditional microstructures. Suitable heat treatments are needed to achieve desired properties.HY100 components were subjected to direct tempering and full heat treatment involving austenize-quench and temper. Heat treatments were conducted at different tempering temperatures and were examined for microstructures and mechanical properties.The results from the tensile tests indicate that direct tempering of HY100 steel can provide good part mechanical properties, but with reduced ductility. A separate austenize /quench step can be eliminated when heat treating SLM HY100 parts.

A separate austenize /quench step can be eliminated when heat treating SLM HY100 parts. However, further experimental studies are recommended to fully understand a proper heat treatment schedule.AcknowledgmentsThe authors gratefully acknowledge the support of the Tim Gornet, Mathew Taylor and Zach Daniel of the Rapid Prototyping Center at the University of Louisville for their help and assistance.References[1] I. Gibson, D.W. Rosen, B. Stucker, Additive Manufacturing Technologies: Rapid Prototyping to Direct Digital Manufacturing, Springer, New York, NY, 2009.[2] K.Kempena, E.Yasa, L.Thijs, J.P. Kruth, J.Van Humbeeck, Microstructure and mechanical properties of Selective Laser Melted 18Ni-300 steel. Physics Procedia 12, 2011, 255–263.[3] J.P. Kruth, L. Froyen, J. Van Vaerenbergh, P. Mercelis, M. Rombouts, B.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperTHE EFFECT OF PROCESS PARAMETERS AND MECHANICAL PROPERTIES OF DIRECT ENERGY DEPOSITED STAINLESS STEEL 316Mojtaba Izadi*, Aidin Farzaneh*, Ian Gibson*, Bernard Rolfe*School of Engineering, Deakin University, Geelong, Victoria, AustraliaAbstractProcess parameters in Direct Energy Deposition (DED) Additive Manufacturing are playing an important role in order to fabricate desired parts. In this research, we studied the effect of 3 process parameters, namely laser power, scan speed and powder feed rate. Based on variation of these parameters, we examined macrostructure and mechanical properties of stainless steel 316 fabricated parts, employing an orthogonal L9 array using the Taguchi technique.

Based on variation of these parameters, we examined macrostructure and mechanical properties of stainless steel 316 fabricated parts, employing an orthogonal L9 array using the Taguchi technique. The results showed laser power to be the most effective factor whereas scan speed and powder feed rate were respectively less effective. In addition, effect of height of deposition was also considered. The results indicated change in macrostructure with increasing height. Finally, validation of a previously defined energy density equation for the DED process was studied. The results clearly showed the current energy density equation cannot fully represent a relation between input energy and output geometry, macrostructure, and mechanical properties.IntroductionAdditive Manufacturing (AM) is a relatively new method of fabrication that is being implemented widely in automotive, medical, aerospace industries [1].

Metal-based AM is increasing at a particularly high rate, focusing on direct manufacture of complex geometry metal parts. Control of process parameters in order to achieve desired mechanical property, macrostructure, microstructure, and dimensional accuracy has been always a challenge on Direct Energy Deposition metal AM processes in particular [2]. Laser Engineered Net Shaping (LENSTM) is a typical process of Direct Energy Deposition and the subject of this research. In this specific process, there are up to four nozzles that deposit powders and a laser which melts this powder to create a melt pool in order to fabricate or repair a metallic part on a substrate [3].To address these problems, firstly it is necessary to have an understanding of the effect of parameters on part properties. In order to reach that goal, Kobryn et al studied the effect of laser power, scan speed, and substrate thickness on porosity for Ti-6Al-4V (Ti6-4) deposits.

In order to reach that goal, Kobryn et al studied the effect of laser power, scan speed, and substrate thickness on porosity for Ti-6Al-4V (Ti6-4) deposits. They found porosity declines with increasing laser power [4]. Krishna et al fabricated several cylindrical titanium samples with different dimensions, and different values of open and close pores with random process parameters such as laser power, scan speed, powder feed rate and hatch distance. They reported compressive strength changes by changing the mentioned process parameters [5]. With the same methodology, Bandyopadhyay et al deposited several samples with different materials such as pure Ti and TiNi alloy and then compared compressive stiffness of samples that have similar porosity[6]. They showed compressive stiffness for NiTi samples are much lower. The main drawback of these studies is that all process parameters were chosen randomly.

They showed compressive stiffness for NiTi samples are much lower. The main drawback of these studies is that all process parameters were chosen randomly. Two further studies, one using pure titanium and one using Ti6-4, were carried out on the effect of porosity distribution on mechanical properties. The result clearly indicated pore distribution is very important on mechanical property for both materials [7, 8].1058In order to overcome the problem of random parameter selection, Krishna et al, used a systematic method to fabricate NiTi cylindrical samples, studying each process parameter individually. They varied one parameter and kept the rest constant. This raises a question, what is the effect of a combination of parameter variations on porosity? To tackle this issue, they employed total energy input per volume of each scan (E), called energy density equation, which is based on Powder Based Systems (PBS).

PBS researchers have found a reasonable relation between energy density and mechanical properties such as density, hardness, wear resistance as well as resulting macrostructure [9-12]. This energy density equation includes laser power (P), scan speed (𝑣), hatch distance or scan line spacing (h) and layer thickness (t) to unify these parameters [13] ;⁄𝐸 = 𝑃 𝑣 × ℎ × 𝑡(1)Ferguson et al. investigated scan speed and powder feed rate while laser power was constant to create an empirical solution. They presented a new unification unit named linear mass density, 𝑑𝑚 𝑑𝑙⁄ (g/mm), which includes laser scan speed and powder feed rate. The aim of this study was to obtain the closest parameter combination to create a non-porous bulk structure over each 𝑐𝑚3of stainless steel 420 or steel 4140.

The aim of this study was to obtain the closest parameter combination to create a non-porous bulk structure over each 𝑐𝑚3of stainless steel 420 or steel 4140. However, in the mentioned study, mechanical properties were not analysed[14].There are also numerous studies investigating the effect of several random process parameters on porosity for different materials by depositing cylindrical samples and machining them to make standard fatigue or tensile samples [15-17]. The main disadvantages of this approach are that there is no guarantee of a homogeneous distribution of porosity on these samples. Porosity is likely to vary in such samples due to uneven heat distribution from the bottom to top. As explored by Manvatkar et al. by increasing the number of layers in a thin wall made of stainless steel 316, wall cooling rate decreases; whereas maximum temperature greatly increases. Microhardness tests in this study showed hardness decreases by increasing height [18].

Microhardness tests in this study showed hardness decreases by increasing height [18]. Interestingly Amine et al. indicated value of hardness in lower layers is higher than the upper layers for Stainless steel 316L in multilayers of thin wall, in another word, hardness decreases at increasing heights [19]. However, Yan et al. found different results for Ti6-4. They indicated that hardness increases as you measure higher up the part [20].In this paper, firstly the relationship between energy density and outputs such as geometry, porosity, and mechanical properties will be studied. Secondly, effect of height on macrostructure also will be investigated. Finally, effect of porosity on mechanical properties will be explored.Material and MethodsIn this study, experimental design based on L9 orthogonal array of Taguchi method is used and three process parameters in three levels; low, mid and high, are investigated. Each level is constant in 3 samples while the other 2 are changed.

Each level is constant in 3 samples while the other 2 are changed. As this study is done to find any potential effect of process parameters on porosity, there was no repetition to validate the results and further study should be done with more samples to calculate the variance. Stainless steel 316 with spherical particle size between 50 to 150 µm was deposited on a substrate of the same material and dimensions of 125mm×125mm×25mm. A Laser Engineered Net Shaping (LENSTM) MR7 (Optomec Inc. Albuquerque, NM) machine was used to build the samples. Input layer thickness and hatch distance were 0.254 and 0.381 mm respectively. In addition, laser focal distance was 9.652 mm. All nine cylindrical specimens were manufactured in a glove box containing Argon atmosphere with 𝑂2 content less than 10 ppm to minimize oxidation during the deposition process.

All nine cylindrical specimens were manufactured in a glove box containing Argon atmosphere with 𝑂2 content less than 10 ppm to minimize oxidation during the deposition process. It should be mentioned that each cylinder was1059fabricated individually, and after each fabrication there was an approximate 10 minutes pause time to let the substrate cool down to normal temperature.Process parameter variations can be seen from Table 1. After part fabrication, dimensions for all parts, by employing micro-scope and Vernier calipers, were measured. Then porosity for the top and bottom each sample was measured using ImageJ 2D software, following removal of the top material to maintain an equal sample height. Finally, compression test in order to measure compressive stiffness was carried out using Instron 50KN.Table1.

Finally, compression test in order to measure compressive stiffness was carried out using Instron 50KN.Table1. Process parameters and energy densitySpecimenLaser power(W)Scan speed(mm/s)Powder feed rate (gram/min)Energy density (J/mm2)L1 L2 L3 L4 L5 L6 L7 L8 L9250 250 250 200 200 200 150 150 1508.5 12.7 17.0 8.5 12.7 17.0 8.5 12.7 17.57.0 10.0 13.5 10.0 13.5 7.0 13.5 7.0 10.0305.12 203.41 152.55 244.09 162.73 122.05 183.07 122.05 91.54123456789Figure1. Samples as depositedResults and discussionsIn this section, the relationship between energy density and dimension, porosity and mechanical properties are studied. Then, the effect of height on porosity will be investigated. Finally, the effect of process parameters on compressive stiffness will be considered.Energy Density and DimensionDifferent process parameters produce different results in the final cylinders’ dimensions. The initial target for dimensions was 6.5mm Ø and 14mm length.

The initial target for dimensions was 6.5mm Ø and 14mm length. These dimensions were chosen so that we could trim the samples by using wire cut, and make them 13mm length as a scaled dimension according to ASTM Standard E9-09 [21]. However, none of the samples were fabricated with the targeted dimensions. Part fabrication dimensions are shown in Table 2. This low accuracy seems due to two reasons. One could be related to process parameters unable to create a proper melt pool. Another reason may be a cumulative effect1060caused by the build strategy used where the outline of the part is deposited, followed by a fill pattern. This issue can be seen more on L6 and L8 (Figure 1), where lack of powder infill is significant.As can be seen from Table 2, there is no specific trend between energy density (as computed using Equation 1) and dimension.

For instance, in L1 and L4, average length increases by decreasing energy density, while in L7 and L9 average length increases by increasing energy density. More interestingly L6 and L8, which have exactly the same energy density, average length is very different (see Figure 2A). Also, the diameter of cylinders is slightly different from bottom and top. Similarly, there is no logical relationship between average diameter of top and bottom and energy density (see Figure 2B). In both figures, volumes are normalized between 0-1 in order to show correlation between energy density and dimensions.Table2.

A) Energy density vs. average length B) Energy density vs. average diameterEnergy density and PorosityIn DED there are several types of pores. Lack of fusion porosity; caused by insufficient/incomplete melting which highly depends on input parameters. Gas porosity; caused by entrapment of gas from the powder feed system and/or gas released from the powder particles. Finally, hatch distance porosity; structure which is created by increasing hatch spacing between two cross-sectional layers [22, 23].Volume of porosity is different in each of the build samples, but there is no apparent relationship between energy density and amount of porosity as seen in Table 3 and Figure 3.1061Sample NumberTable3.

Energy density and average porositySpecimenEnergy density (J/mm2)Ave Porosity (%)L1305.121.18L2203.411.77L3152.552.10L4244.092.98L5162.735.96L6122.056.67L7183.0715.49L8122.0515.11L991.5415.82MAXMINSample NumberFigure 3. Relation between energy density and average porosityEnergy Density and Mechanical PropertiesWe also studied the relationship between the energy density and compressive stiffness. In this case, although a trend between energy densities of samples and compressive stiffness can be seen, energy density figures could not be used to determine the exact compressive stiffness (see Table 4 and Figure 4).Table4. Energy Density and mechanical propertiesSpecimenL1Energy density (J/mm2)305.12Compressive stiffness (GPa ) 520L2203.41310L3152.55271L4244.09380L5162.72120L6122.05124L7183.07118L8122.0571L991.54631062MAXMINSample NumberFigure 4.

Relationship between Energy density and Mechanical propertiesMechanical Properties and PorosityIn order to understand the effect of porosity on mechanical property, compression tests were carried out on the samples. The sample sizes were not standard according to ASTM, which may have resulted in a measurement error in L1 which has an unusual 520GPa value for stiffness. However, it is considered that they still can be compared against each other. The error may be due to the close strain control points during the test and so could be avoided by using longer specimens in future. The trend of results clearly shows by increasing porosity, compressive stiffness decreases. However, there are two exceptions. Specimen number 4 has higher porosity compared with L2 and L3, but it also has higher compressive stiffness. In addition, specimen number 7 has higher porosity than L8 and also has higher compressive stiffness. This could be the effect of grain sizes or pore distribution.

In addition, specimen number 7 has higher porosity than L8 and also has higher compressive stiffness. This could be the effect of grain sizes or pore distribution. These assumptions should be validated in future by microscopic evaluation of grain size and the more comprehensive use of ImageJ 2D to find the pore distributions.MAXMINSample numberFigure 5. Relation between and average porosity volume and compressive stiffnessTable 5. Relation between Porosity volume and Compressive stiffness (Gp)SpecimenL1 L2 L3 L4 L5 L6 L7 L8 L9Porosity on bottom (%) 1.85 2.61 2.57 3.54 8.05 6.52 15.87 17.12 17.92Porosity on top (%)0.51 0.92 1.62 2.41 3.86 6.81 15.11 13.10 13.71Average porosity (%) 1.18 1.77 2.10 2.98 5.96 6.67 15.49 15.11 15.82Compressive Stiffness (GPa) 520 310 271 380 120 124 118 71 631063Height and PorosityIn deposited samples, it is seen that there is more porosity on the bottom than near the top in most cases (see Figure 6 and 7).

This is likely due to high cooling rate on layers which are close to the substrate (see Table 5). However, sample number L6 is an exception. This sample was observed to be the shortest sample and therefore possibly there was insufficient powder for each layer and the laser was therefore possibly remelting previous layers and thus increasing density (see Figure 8).L1 L2 L3 L4 L5L6 L7 L8 L9Figure 6. Bottom of specimensL1 L2 L3 L4 L5L6 L7 L8 L9Figure 7. Top of the specimen10641 mm1 mmPorosity 20.00%18.00%16.00%14.00%12.00%10.00%8.00%6.00%4.00%2.00%0.00%L1L2L3L4L5L6L7L8Porosity on bottomPorosity on topFigure 8. Porosity on bottom and topEffect of process parameters on porosityDifferent process parameters have different effects on porosity. Laser power by rank 1 has the largest effect on porosity. Scan speed and powder feed rate stand on second and third rank respectively (See Table 6).

Laser power by rank 1 has the largest effect on porosity. Scan speed and powder feed rate stand on second and third rank respectively (See Table 6). These ranks are obtained through Minitab statistical analysis software.Table6. Ranking effect of process parameter on bottom porosityLevel 1 2 3 RankLaser power 16.972 6.038 2.344 1Scan speed 7.086 9.262 9.006 2Feed rate 8.498 8.025 8.831 3DOE Results validationThe means for each level are considered as the average of the response achieved at the same control factor levels in a static Taguchi DOE (see Table 6). The analysis of means of the experiment show which parameters have more influence on density.S/N ratios evaluate the effect of noise, which is a measure of the process robustness, of the response to variability of a parameter’s influence.

The general nature of this graph is the same as the means effect (see Figure 9).In order to check the validity of the results from the DOE, mean effect graph for means and signal to noise (S/N) ratio should be considered. The lowest values for each factor in Means plots should be equal to the highest value of S/N ratio which can be seen in fig 9. Hence, these graphs indicate the DOE is validated.1065L9Sample NumberFigure 9. Means vs. SN ration for bottom porosityConclusionBy increasing the height, the amount of porosity in general decreases, probably due to decreasing cooling rate on top layers. This also affects mechanical properties. Hence, there is a need to control the macrostructure, and consequently mechanical properties, in order to achieve a homogeneous structure. In addition, according to the results, there is difficulty in finding a correlation between energy density and outputs. It is likely that the energy density equation is incorrect for DED.

In addition, according to the results, there is difficulty in finding a correlation between energy density and outputs. It is likely that the energy density equation is incorrect for DED. This could be improved by including parameters relating to powder feed rate, thermal transfer properties, and height above the substrate, etc. These results certainly indicate that energy density calculation in DED is more complex than for powder bed, and perhaps could be replaced by a function that presents a logical correlation between input parameters and outputs.References[1][2][3][4][5][6][7][8]I. Campbell, D. Bourell, and I. Gibson, "Additive manufacturing: rapid prototyping comes of age," Rapid Prototyping Journal, vol. 18, pp. 255-258, 2012. G. Sun, R. Zhou, J. Lu, and J. Mazumder, "Evaluation of defect density, microstructure, residual stress, elastic modulus, hardness and strength of laser-deposited AISI 4340 steel," Acta Materialia", vol. 84, pp. 172-189, 2015. Y.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of Process Parameters and Shot Peening on Mechanical Behavior of ABS Parts Manufactured by Fused Filament Fabrication (FFF)Cody Kanger, Haitham Hadidi, Sneha Akula, Chandler Sandman, Jacob Quint, Mahdi Alsunni, Ryan Underwood, Cody Slafter, Jason Sonderup, Mason Spilinek, John Casias, P. Rao, M.P. Sealy1Dept. of Mechanical and Materials Eng., University of Nebraska, Lincoln, NE 68588, U.S.A.AbstractThe goal of this research was to understand how shot peening affected the tensile strength and elongation of ABS polymer parts between three process parameters: layer height, infill angle, and outer shell quantity. Experiments were conducted using a Hyrel 30M fused filament fabrication (FFF) printer to produce ASTM 638D-IV samples.

Experiments were conducted using a Hyrel 30M fused filament fabrication (FFF) printer to produce ASTM 638D-IV samples. This is an important area of research because 3D printed polymers have typically been limited to prototyping applications due to low strengths and stiffness. Traditional means of improving a polymer’s mechanical properties are changing the structural or chemical makeup. However, shot peening, a surface treatment commonly used to improve mechanical properties of metals, was hypothesized to have a statistically significant effect on the tensile strength and elongation of polymer parts. Results showed that shot peening had a significant effect on decreasing the tensile strength. Although not statistically significant, samples did show an increase in elongation after shot peening.Keywords: fused filament fabrication, shot peening, ABS, strength, elongation1.

Although not statistically significant, samples did show an increase in elongation after shot peening.Keywords: fused filament fabrication, shot peening, ABS, strength, elongation1. IntroductionThree-dimensional printing is a manufacturing practice that is reshaping industries and producing new opportunities. Polymer printing is a widely studied area in additive manufacturing as it has vast capabilities and is an affordable manufacturing method for custom parts. However, depending on 3D printer quality, mechanical properties vary widely. As desktop printers become more commonplace with hobbyist, small business owners, and K-12 education, the need to print low cost, high quality components quickly and cheaply becomes challenging. There are several methods to improve the mechanical properties of polymer printed parts. The first approach to achieving higher quality parts is optimizing print process parameters.

There are several methods to improve the mechanical properties of polymer printed parts. The first approach to achieving higher quality parts is optimizing print process parameters. Another option is post-processing that uses chemical, physical, or thermal mechanism to refine mechanical properties. A lesser explored option is the use of surface treatments, such as shot peening, to improve the surface mechanical properties of polymer parts. The use of surface treatments to improve part quality on low-cost printers may be an effective approach to achieve similar results seen by high-end 3D printers.1 Corresponding author: sealy@unl.edu (M. Sealy)444Polymer printing has copious avenues of exploration relationship between process parameters and part quality. A common printing process filament fabrication (FFF); see Fig. 1. FFF has numerous process parameters leaving voids in the knowledge about their effects on part quality.

A common printing process filament fabrication (FFF); see Fig. 1. FFF has numerous process parameters leaving voids in the knowledge about their effects on part quality. The FFF process consists of a printer head extruding material (usually heated) onto a build platform producing a 2-dimensional design. This design comes from a 3D model that is sliced into layers and collectively results in a 3D part. FFF allows for nearly any design of a part without additional tooling. Entry-level printers often produce parts with inferior mechanical properties to those produced by traditional processes, such as injection molding. The strength of FFF polymer parts is dependent on parameters such as:Extruder Head (235°C)to understandingtheisfusedFilament tracksFilamentTape SubstratePrint Bed (85°C)Fig. 1 Fused Filament Fabrication (FFF).2D design layer thickness, Fig. 2(a)  Polymer track orientation, thickness, and spacing, Fig. 2(b)  No.

For example, it was shown that increasing the number of outer shells in a part increases the tensile strength, but reduces elongation characteristics [1-4]. Also, data has shown that smaller layer heights resulted in the highest tensile strength due to an increased cross sectional area according to Garg and Bhattacharya [5]. However, the next strongest sample had the largest layer height tested and was attributed to a reduction in air pockets where cracks would initiate [5, 6]. Raster angle also has variability in research results. Some sources printed consistent raster angles throughout, while others alternated directions between layers. Most sources agreed [5-7] that a 0° raster angle (parallel to loading) resulted in the best tensile strength when maintained throughout the part. Additional research revealed that when alternating between build layers, a 0°/90° setup was better compared to only 0°, 45°, or 90°445angles [8, 9].

Additional research revealed that when alternating between build layers, a 0°/90° setup was better compared to only 0°, 45°, or 90°445angles [8, 9]. Even still, others have found that 60° and higher raster angles resulted in a stronger tensile strength [10, 11]. This indicates inconsistent results on raster angle’s effect on strength.Post processing methods are another way to manipulate material properties like tensile strength and elongation. Many methods exist, but shot peening is a quick, low cost, and straight-forward option. Shot peening uses small spheres (media) at a high velocity to plasticize a surface, inducing a beneficial thus compressive residual stress, (Fig. 3). This stress increases crack resistance in the outer surface of the material and alters other mechanical properties [12]. Shot peening parameters can be adjusted depending on the user’s goal.

3). This stress increases crack resistance in the outer surface of the material and alters other mechanical properties [12]. Shot peening parameters can be adjusted depending on the user’s goal. First, the media size and material can be changed depending on the amount of surface area to cover and what type of surface is present. Peening speed, which is dictated by pressure and nozzle diameter, can change to increase the depth of penetration [13].Compressed AirNozzleShot Peening MediaResidual StressWorkpiece(-)(+)Fig. 3 Shot peening process.Both FFF and shot peening are useful processes that allow many parameters to be adjusted. There was no reported literature on the use of shot peening on polymers, most likely due to the limited structural use of polymers. All reported shot peening literature related to additive manufacturing has been conducted on metal alloys.

All reported shot peening literature related to additive manufacturing has been conducted on metal alloys. The purpose of this research was to investigate the strength and elongation of an ABS polymer printed on an entry-level machine by adjusting process parameters and incorporating shot peening. Shot peening was implemented due to its effectiveness of changing mechanical properties and efficient cost of operation. Results from this experiment will determine the feasibility to use shot peening as a supplemental processing step to improve mechanical properties.2. Experimental Setup2.1 Fused Filament FabricationTwo different types of samples were printed using ABS filament with a 1.75 mm diameter. First, an ASTM 638D Type IV “dog bone” part shown in Fig. 4(a) was used for the purpose of tensile and fracture strength testing in addition to elongation measurements. Second, a flat pseudo-Almen strip sample as shown in Fig.

4(a) was used for the purpose of tensile and fracture strength testing in addition to elongation measurements. Second, a flat pseudo-Almen strip sample as shown in Fig. 4(b) was dimensionally identical to an A-2 Almen strip and used for preliminary work to determine shot peening process parameters.(a)(b)20 mmFig. 4 Sample geometry: (a) ASTM 638D Type IV and (b) pseudo-Almen strips.446For both samples, three print process parameters were changed as part of a central composite design to maximize efficiency. These parameters had low, center, and high values which were within the acceptable range of each parameter’s capability. Table 1 shows the levels for layer height, raster angle, and number of outer shells used in this experiment. These values produced nine total “printing recipes” which consisted of (a) eight recipes with every combination of the low and high variables and (b) one recipe exclusively with the center points of each variable.

Each of the eight recipes of low/high had 3 repetitions (n = 3), while the center point had 9 repetitions (n = 9). In addition to changing the main process parameters, the material flow rate factor (dependent process parameter) was adjusted to accurately achieve the desired layer heights, as seen in Fig. 5. All other printing parameters can be referenced in Table 3A in the Appendix.Table 1 Levels of Layer height, No. of Outer Shells, and Raster AnglesVariable Layer Height No. of outer Shells (X-Y plane) 2,3,4 Raster Angle (from loading)Levels (Low, center, high) 0.12 mm, 0.26 mm, 0.40 mm9°,23°,36°100 µmGauge length cross section view100 µmFig. 5 Microscopic image of layer height: (a) low layer height (0.12 mm) and (b) high layer height (0.40 mm).All parts were printed on a Hyrel 30M fused filament fabrication (FFF) machine as seen in Fig. 6. A specific procedure was executed during the production of parts, which included part placement, printing, and removal.

6. A specific procedure was executed during the production of parts, which included part placement, printing, and removal. Only three parts were printed at one time, as shown in Fig. 7, in order to be efficient while reducing random thermal gradients. Manufacturing of the parts started when the minimum print bed and nozzle temperatures were met (Table 3A in the appendix). Prior to all print jobs, the machine’s z-calibration, which is the distance between the bed and nozzle, was either validated or recalibrated to be at the manufacturer’s recommended distance (approximately 1 mm). Finally, parts were removed using a razor blade to gently “pop” the longer447edge of the part below the painters’ tape substrate as shown in Fig. 7. The tape allowed for better adhesion to the print bed and a buffer for removing parts to avoid damage with the razor blade.ABS 1.75 mm filamentExtruder HeadZYPrint BedSamplesXFig. 6 Hyrel 30M fused depositionFig.

6 Hyrel 30M fused depositionFig. 7 FFF build platform.modeling (FDM) machine.2.2 Shot PeeningShot peening severely bent the samples due to imbedded residual stresses. Here, the pseudo-Almen strip samples were shot peened to exploit their ability to easily measure the curvature resulting after peening. Pseudo refers to the fact that Almen strips pertain strictly to metals to determine saturation. The change in curvature, or deflection, was measured using pictures taken immediately after peening of the sample parallel to the peened surface. The deflection of each strip was measured in seven locations equally spaced from the center (Fig. 8). With this procedure, one combination of variables (or recipe) was peened with 10, 20, 30, 40, or 50 passes at a rate of 10 passes per minute (3 samples for each recipes). The deflected strips were used to develop the saturation curve, which was later validated with all remaining recipes by comparing their deflections for consistency.

The deflected strips were used to develop the saturation curve, which was later validated with all remaining recipes by comparing their deflections for consistency. At 20 passes, all recipes had negligible differences in their deflection indicating a saturation point where significant further deflection was minimal. Table 2 gives the parameters used for shot peening the dog bone samples. Shot peening was done using a Westward blast cabinet equipment with a pressure regulator. To produce consistent results, the jig in Fig. 9 was used to secure the sample.Reference Line10 mmFig. 8 Pseduo-Almen strip deflection after shot peening.448Pseudo-Almen stripTable 2 Shot Peening Process ParametersValue Parameter Glass Media Type 0.4-0.6 mm Media Size 0.55 MPa Air Pressure Nozzle Diameter 4.76 mm Contact Distance 100 mm No. of passes Application Rate 10 passes/minute20Peening nozzleFig.

of passes Application Rate 10 passes/minute20Peening nozzleFig. 9 Shot peening fixture.2.3 Tensile TestingTensile testing was performed on a MTS 810 equipped with a 25-kN capacity load-cell and an extensometer measuring a distance of 25 mm as seen in Fig. 10. The non-gauge length ends of the samples were securely fastened in the upper and lower grips. The strain rate was 0.01 mm/min and the data sampling rate was 0.10 seconds. Due to a testing failure, only 8 were used in the results. Calipers were used prior to fracturing to measure the actual sample gauge width and height. With the cross-sectional area, stress for each sample was calculated.Applied loadSampleExtensometer SampleGripsFig. 10 MTS 810 tensile testing machine setup.4493. Results and Discussion3.1 Elongation3.1.1 Scatter Plot ResultsLayer height and shot peening were found to influence elongation (Fig. 11(a)).

10 MTS 810 tensile testing machine setup.4493. Results and Discussion3.1 Elongation3.1.1 Scatter Plot ResultsLayer height and shot peening were found to influence elongation (Fig. 11(a)). The peening process increased the 0.12 mm and 0.40 mm layer heights’ maximum elongation from 2.34 mm to 2.50 mm and 2.35 mm to 2.42 mm, respectively. However, shot peening decreased the elongation effects for the 0.26 mm layer height from 2.13 mm to 2.10 mm. The smaller layer height accepted the shot peening process as beneficial because the inside material was well packed (lots of small voids between raster tracks). In contrast, the larger layer height beneficially accepted the surface treatments because its inner material was also well packed, but due to fewer large air voids between raster tracks. The center point of the central composite design hits a critical point where numerous sizeable air voids occur between raster tracks and dilute the effects of shot peening.

The center point of the central composite design hits a critical point where numerous sizeable air voids occur between raster tracks and dilute the effects of shot peening. This same logic of air voids between layers is the reason why the 0.12 mm and 0.40 mm layer heights performed similarly. In contrast, the 0.26 mm layer height resulted in the lowest elongation because its cross sectional area was the least due to a poor balance of thick raster tracks and voids between rasters.4)m m(n o3i ta g n o El2. x a M10.10.20.30.4918273623(a)Layer Height (mm)(b)Fill Angle (°)(c)No. of Outer ShellsFig. 11 Effect of print parameters and shot peening on elongation: (a) layer height, (b) fill angle, and (c) No. of outer shells.Raster angle’s most significant contribution with the addition of shot peening was increasing the 9° and 36° maximum measured elongation.

of outer shells.Raster angle’s most significant contribution with the addition of shot peening was increasing the 9° and 36° maximum measured elongation. Both raster angles (9° and 36°) experienced increased elongation effects with shot peening from 2.33 mm to 2.42 mm and 2.36 mm to 2.49 mm, respectively. At the center point of 22.5°, shot peening decreased the elongation of samples from 2.13 mm to 2.10 mm. When looking at the scatter plot in Fig. 11(b), the range of measured values for the 9° infill angle is vast, but has a concentrated cluster around the 2 mm mark. With this dense area of data, and a handful of outliers that measured 3.5 mm or greater elongation, these4504outliers pulled the average of the 9° samples above that of the 22.5° raster angle samples.

If the data had not followed this path, it would have shown an increased infill angle (up to 45°) would increase elongation, thus agreeing with prior literature.Samples with two and three outer shells had a negligible effect on elongation, i.e., less than a 2% change. When printing at these conditions, on average, samples differed in elongation trends with values changing from 2.10 mm to 2.13 mm and 2.13 mm to 2.10 mm with shot peening, respectively. With 4 outer shells, shot peening had a noticeable effect with increased elongation from 2.59 mm to 2.79 mm after peening. This 10% increase in elongation is believed to occur because at 4 outer shells, enough peripheral material was present for shot peening to beneficially manipulate the mechanical properties of this section and reduce crack initiation and growth. Fig.

Fig. 11(c) shows that more outer shells increased the elongation capabilities for the same reason previously stated; i.e., more outer material slows crack propagation.3.1.2 Main Effects and InteractionsAn effects plot that includes interaction terms for shot peened and non-shot peened ABS tensile samples is shown in Fig. 12 to convey what parameters were most influential. Shot peening increased the displacement for the majority of printed parts. This figure shows adjusting the quantity of outer shells alone had the greatest effect on elongation. Adjusting the layer height alone had the least significant effect on elongation. For most samples, adjusting any two parameters or all three at once had a greater influence over elongation than individual parameters. Finally, these results were consistent with peened and non-peened samples except for No.

Finally, these results were consistent with peened and non-peened samples except for No. of shells and infill angle, which was the third most influential factor over all three parameters when not being peened.)m m(n o10.8Shot PeenedNon-Shot Peenedi ta g n o E n ol0.60.4t c e0.2f fE0Process ParameterLH: Layer Height NS: No. of Outer Shells IA: Raster AngleFig. 12 Main and interaction effects of process parameters and shot peening on elongation.Due to the small amount of change in the elongation of all samples, an ANOVA test was conducted to determine the significance of each of the three process parameters, plus the peening451condition. The test was conducted using a 95% confidence level. Table 3 lists the results of this ANOVA test for all four process parameters. Infill angle was the only parameter which had elongation results that showed significance. The other two printing parameters and the peening parameter were not significant.

Infill angle was the only parameter which had elongation results that showed significance. The other two printing parameters and the peening parameter were not significant. That is, the p-values were greater than 0.05.Table 3 ANOVA test with 95% confidence level for displacementVariable Layer Height Infill Angle No. Outer Shells Shot-PeeningP-Value 0.811 0.000 0.729 0.4523.2 Tensile and Fracture Strengths3.2.1 Scatter Plot ResultsTensile and fracture strength results for various layer heights had similar trends regarding printing process parameters and shot peening effects. After shot peening, tensile strength changed from 26 MPa to 24 MPa and 22 MPa to 23 MPa for the 0.12 mm and 0.26 mm layer heights, respectively (Fig. 13). With a layer height of 0.40 mm, shot peening had no effect as non-peened and peened samples averaged 24 MPa of tensile strength.

13). With a layer height of 0.40 mm, shot peening had no effect as non-peened and peened samples averaged 24 MPa of tensile strength. Ultimate fracture strengths for 0.12 mm and 0.26 mm samples resulted in changes of 24 MPa to 21 MPa and 20 MPa to 21 MPa after shot peening. Similar to ultimate tensile strengths, the fracture strength did not change, on average, after shot peening as both non-peened and peened samples resulted has a fracture strength of 22 MPa. Consistently, the 0.26 mm layer height resulted in the lowest tensile and fracture strengths with or without peening. The maximum tensile and fracture strengths for non-peened samples were produced at the 0.12 mm layer height in contrast to peened samples, which were at the 0.40 mm. This is because the smaller layer heights created more surface area within the cross sectional area.

This is because the smaller layer heights created more surface area within the cross sectional area. But, with shot peening, these thin layer heights were damaged and could not collect residual stress, thus leading to a decline in their respective tensile and fracture strengths.Outer shell quantity had similar trends for ultimate tensile and fracture strengths when shot peening was introduced (Fig. 14). For two outer shells, a decrease in tensile strength from 21 MPa to 20 MPa was observed for samples which were peened. For fracture strength, no change occurred, on average, as both non-peened and peened samples has a fracture strength of 19 MPa. At three outer shells, shot peening increased both the tensile and fracture stresses by about 5%. Non-peened samples exhibited 22 MPa of tensile strength, and peening them increased that to 23 MPa. The opposite occurred with four outer shells.

Non-peened samples exhibited 22 MPa of tensile strength, and peening them increased that to 23 MPa. The opposite occurred with four outer shells. Non-peened samples exhibited a tensile strength of 29 MPa compared to peened samples with a tensile strength of 28 MPa. Fracture strength of non-peened samples measured 26 MPa while peened samples were 25 MPa. Therefore, at four outer shells, about a 3.5% decrease in both tensile and fracture strengths resulted. The quantity of outer shells as a printing process had a more significant effect on ultimate tensile and fracture strengths as both increased by approximately 35% for non-peened and peened samples.452htg n e r tS el is n e T x a M) a P M(35302520htg n e r tS e r ut c a r F x a M) a P M(3530252015150.10.20.30.40.10.20.30.4(a)Layer Height (mm)(b)Layer Height (mm)Fig.

13 Effect of layer height and shot peening on (a) tensile strength and (b) fracture strength.htg n e r tS el is n e T x a M) a P M(35302520htg n e r tS e r ut c a r F x a M) a P M(3530252015151.822.833.841.822.833.84(a)No. Outer Shells(b)No. Outer ShellsFig. 14 Effect of outer shells and shot peening on (a) tensile strength and (b) fracture strength.Raster angle exhibited consistent trends between printing parameters and peened/non-peened samples (Fig. 15). Ultimate tensile strength changed for both the 9° and 22.5° raster angles going from 25 MPa to 23 MPa and 22 MPa to 23 MPa, respectively. The fracture strength also changed in a similar manner, 23 MPa to 21 MPa, and 20 MPa to 21 MPa for 9° and 22.5°, respectively. For the 36° samples, no change occurred as the tensile strength was 25 MPa and the fracture strength was 23 MPa.

For the 36° samples, no change occurred as the tensile strength was 25 MPa and the fracture strength was 23 MPa. This trend was consistent with literature; that is, the 45° alternating raster angle exhibits the greatest tensile and fracture strength as compared to 0°, 90°, or 0°/90° patterns.453ht35ht35g n e r tS el is n e T x a M) a P M(302520g n e r tS e r ut c a r F x a M) a P M(302520151551525355152535(a)Fill Angle (°)(b)Fill Angle (°)Fig. 15 Effect of fill angle and shot peening on (a) tensile strength and (b) fracture strength.3.2.2 Main Effects and InteractionsFig. 16 plots the effectiveness of each individual parameter and the interaction terms to show which are most sensitive to changing the tensile strength. For both non-peened and peened samples, No. of outer shells had the greatest effect on manipulating the tensile strength of a printed part.

For both non-peened and peened samples, No. of outer shells had the greatest effect on manipulating the tensile strength of a printed part. This was because outer shell quantity dictates how much material through which a crack must propagate. Without peening, layer height was second for the most influential parameter regarding tensile strength. However, it was nearly negligible after peening. With and without peening, adjusting both infill angle and outer shell quantity had the least effect on tensile strength except for shot peened samples where all three parameters were changed.htg n e r tS el is n e T n o108) a P M6(4Shot Peened Non-Shot Peenedt c e2f fE0Process ParameterLH: Layer Height NS: No. of Outer Shells IA: Raster AngleFig. 16 Main and interaction effects of process parameters and shot peening on tensile strength.454To build upon the results of this section, an ANOVA test was conducted on the tensile strength data.

16 Main and interaction effects of process parameters and shot peening on tensile strength.454To build upon the results of this section, an ANOVA test was conducted on the tensile strength data. This 95% confidence level test was conducted because similarly to the elongation results, the changes in tensile strength between samples was small relative to their values. However, the results of the ANOVA test shown in Table 4 were significant for almost all fields. Infill angle, number of outer shells, and the shot peening parameters were well within tolerance for being significant for the 95% confidence level. Only the layer height term was not significant, just narrowly being outside the threshold for significance (0.05 in this study).Table 4 ANOVA test with 95% confidence level for tensile testingVariable Layer Height Infill Angle No. Outer Shells Shot-PeeningP-Value 0.071 0.000 0.013 0.0084.

Outer Shells Shot-PeeningP-Value 0.071 0.000 0.013 0.0084. Summary & ConclusionsSeveral trends were observed by adjusting 3D printing process parameters and applying shot peening as a surface treatment to ABS printed parts on a fused filament fabrication (FFF) system. First, Fig. 17 shows a generalized stress-strain schematic of observed trends. In nearly every printing recipe (except one), shot peening reduced the ultimate tensile and fracture strengths, which may be attributed to multiple factors. One factor that may have caused reduced strength was improper shot peening practices. Over peening or peening with worn glass beads (sharp, small fractured beads as compared to smooth circular ones) can damage the surface and even produce micro-cracks instead of inducing favorable compressive residual stresses. Another possible explanation was that polymers do not form beneficial residual stress fields as well as metals.

Another possible explanation was that polymers do not form beneficial residual stress fields as well as metals. The second noticeable trend from this study was that most samples exhibited increased elongation. This trend was more substantial on samples printed with four outer shells. Having more layers on the outer edge led to parts with better elongation characteristics because there was more material for cracks to propagate through. Additionally, this material was able to beneficially absorb some of the shot peening effects and reduce crack propagation.Another important note from this experiment was that the number of outer shells was the most influential factor in determining ultimate tensile strength and maximum elongation, regardless of whether samples were peened or not. This means that when choosing to adjust layer height, raster angle, or No. of outer shells, changing the outer shell quantity was the most effective.

This means that when choosing to adjust layer height, raster angle, or No. of outer shells, changing the outer shell quantity was the most effective. Another trend was that all three processing parameters were more influential in groups then individually when it came to maximum elongation. This was not as observable for ultimate tensile strength, as other printing process combinations were equally as influential and less than outer shell’s effectiveness by a factor of three. In addition, ANOVA tests were conducted on all 4 process parameters (printing and peening) which revealed a decrease in tensile strength with significance (95% confidence level), except when adjusting the layer height. In contrast, only the infill angle parameter had significance with regards to the results for elongation.455s s e r tSNon-Peened PeenedStrainFig.

In contrast, only the infill angle parameter had significance with regards to the results for elongation.455s s e r tSNon-Peened PeenedStrainFig. 17 Representation of stress-strain results for non-peened and shot peened ASTM 638D.These results show the variability of FFF entry-level printers. Without any surface treatment, it was shown that adjusting process parameters for ABS on this FFF Hyrel 30M printer resulted in tensile strength and elongation trends that contradicted prior research. Additionally, shot peening was shown to have an impact on the mechanical behavior of polymers. Although tensile strength decreased after shot peening, maximum elongation was improved. This shows that shot peening was effective as a supplemental solution to optimizing mechanical properties in a way that was more efficient than using a professional grade printer initially.

This shows that shot peening was effective as a supplemental solution to optimizing mechanical properties in a way that was more efficient than using a professional grade printer initially. Both technologies have their place, but this research shows there is multiple ways to achieve a desired result regarding 3D polymer printing and merits further study.5. AcknowledgementsUndergraduate and graduate students conducted this research as a course project in MECH 498/898: Additive Manufacturing at the University of Nebraska-Lincoln (UNL). The authors would like to acknowledge the support of the Department Chair in Mechanical and Materials Engineering at UNL, Dr. Jeffrey Shield, for enabling Drs. Rao and Sealy to develop a research-based course in additive manufacturing in order to spur more undergraduate involvement in research.6. References[1] Farzadi, A., Waran, V., Solati-Hashjin, M., Rahman, Z.A.A., Asadi, M., Osman, N.A.A., 2015.

Shot peening is a surface treatment to improve the mechanical properties of components. The goal of this work was to quantify the effect of SLA print process parameters, namely layer height and UV exposure, and shot peening on the longitudinal tensile strength of ASTM D638 Type 5 test artifacts. Test parts were created using a central composite experimental plan on a B9 Creator desktop SLA machine. Deflection of the pseudo-Almen strips after shot peening was measured using a digital camera to identify desired peening condition. Post-shot peening tensile strength was measured for the ASTM D638 Type 5 parts. Shot peening generally decreased the strength of MIP- SLA parts.Keywords: stereolithography, shot peening, polymer, strength, deflection1. IntroductionThe most common application of 3D printed polymer parts is rapid prototyping. In rapid prototyping, builds allow for design visualization, assembly verification, and functional testing.

IntroductionThe most common application of 3D printed polymer parts is rapid prototyping. In rapid prototyping, builds allow for design visualization, assembly verification, and functional testing. However, these parts are not commonly used for fully functional applications because of limited mechanical properties. In order to use these polymer parts as functional parts, there is a need to have a better understanding of how process parameters affect build quality and mechanical properties as well as a method to improve them.To improve mechanical properties of polymers, post processes like infiltration, cryogenic treatment, electroplating can be used [1-4]. An alternative method is mechanical surface treatments. Surface treatments such as laser shock peening, shot peening, burnishing, and ultrasonic peening are well known for improving properties of a material. This is done by inducing compressive residual stresses in subsurface of parts through plastic deformation.

This is done by inducing compressive residual stresses in subsurface of parts through plastic deformation. Usually, these surface treatments are performed on metals. However, the effect of these surface treatments on a 3D printed polymer is not well known.There are two objectives for this paper: (1) determine the effect of mask image projection stereolithography (MIP-SLA) process parameters on dimensional accuracy and tensile strength of1 Corresponding author: sealy@unl.edu (M. Sealy)1761polymer parts and (2) determine the effect of shot peening on mechanical properties (tensile strength) of parts made by MIP-SLA (Fig. 1). For shot peening these polymers, pseudo-Almen strips were printed on B9 Creator desktop SLA system. Pseudo-Almen strips were used to measure deflection after shot peening to find the saturation point. Saturation is the point at which further peening produces little to no change in the deflection.

Pseudo-Almen strips were used to measure deflection after shot peening to find the saturation point. Saturation is the point at which further peening produces little to no change in the deflection. Once the saturation point is determined, ASTM D638 Type 5 samples were shot peened to find its effect on mechanical properties (i.e., tensile strength) of the photopolymer.Tensile testerShot peened surfaceModified surface integrity++Polymer SampleFig. 1 Shot peening the surface of an MIP-SLA polymer tensile specimen to determine the effect on mechanical behavior.2. StereolithographyStereolithography (SLA) is based on the process of photopolymerization, which makes use of curable resins that react to ultraviolet (UV) radiation. A part’s build time in SLA depends on the size, complexity, and layer thickness. SLA polymer parts need to be post-cured under UV light to have increased stiffness.

A part’s build time in SLA depends on the size, complexity, and layer thickness. SLA polymer parts need to be post-cured under UV light to have increased stiffness. However, this process imposes an unexpected shrinkage and affects the dimensional accuracy. Wang et al. found that this error can be accounted for by manipulating process parameters [5]. Using the least squares method, a strong correlation existed between shrinkage and the process parameters such that a regression model was developed to predict distortion from nominal conditions. Other parameters like layer thickness also influence the mechanical properties of SLA parts [6]. Typically, smaller layer thickness results in a higher strength.In order to achieve the high precision required by some industries, the stereolithography process must be optimized to achieve better quality builds. To this aim, Campanelli et al.

To this aim, Campanelli et al. performed a statistical analysis to specify the optimal parameter settings resulting in the best dimensional accuracy of printed parts [7]. Considering the two printing resolutions, they found that by setting the proper layer thickness, hatch over cure (HO), and border over cure (BO), the1762post processing requirement such as post curing can be neglected as the whole resin will be cured using the optimal printing parameters.in obtaining homogeneous polymerization. Chartier et al. investigated the influence of different process parameters, including laser power, scanning speed, and number of irradiations on the degree of polymerization to reach the lowest internal stresses and risk of deformation [8]. They showed that a low scanning speed along with a limited laser power lead to the more homogeneous polymer but it increases the build time.

They showed that a low scanning speed along with a limited laser power lead to the more homogeneous polymer but it increases the build time. They also revealed that based on the layer thickness, the number of affected layers by irradiation of the upper layers will be changed.Irradiation parameters are also critical3. Shot PeeningShot peening is a mechanical surface treatment where a stochastic stream of particles impinges the surface to impart favorable compressive residual stresses by severe plastic deformation (Fig. 2). The particles, also referred to as peening media, can be composed of metal, glass, or ceramic beads that are on the range of a few hundred microns to several millimeters in diameter. Shot peening is a low cost and quick process to enhance a large surface area component. The shot acts like a hammer, dimpling the surface and causing compressive stresses under the dimple.

Shot peening is a low cost and quick process to enhance a large surface area component. The shot acts like a hammer, dimpling the surface and causing compressive stresses under the dimple. As shots continue to strike the surface, multiple dimples form as compressive stresses continue to build in the workpiece. These stresses decrease the risk of static and fatigue failures. The use of shot peening on polymers is poorly understood since this process is typically reserved for metals and sometimes ceramics. This work investigates the use of shot peening to modify the mechanical behavior of SLA tensile test specimens.Compressed AirNozzleShot Peening MediaResidual StressWorkpiece(-)(+)Fig. 2 Shot peening process schematic.Fig. 3 B9 Creator mask image projection stereolithography (MIP-SLA) printer.17634.

2 Shot peening process schematic.Fig. 3 B9 Creator mask image projection stereolithography (MIP-SLA) printer.17634. Experimental Procedure4.1 Mask Image Projection Stereolithography (MIP-SLA)The SLA machine used for printing parts (pseudo-Almen strips and ASTM tensile samples) was a B9 Creator (Fig. 3) with a printing resolution of 30 µm to 100 µm. Given the flexibility of the printed parts, several resins were tested to determine which one provide the stiffest part. As a result, the Monocure 3D Blue Resin provided the most accurate and stiff parts.In order to evaluate the effect of process parameters on the printed parts’ quality, layer thickness and UV sensitivity were studied. From the literature review above, layer thickness highly influences build quality and mechanical properties of SLA polymers.

From the literature review above, layer thickness highly influences build quality and mechanical properties of SLA polymers. UV sensitivity determines the length of time each layer is exposed to the UV light, which can affect how much each layer is going to be cured.Using a central composite design, cross sectional area (related to the middle part) and tensile strength were measured for ASTM D638 Type 5 samples. The sample geometry is provided in Fig. 4. The statistical parameters used for the experiment are shown in Table 1.

The sample geometry is provided in Fig. 4. The statistical parameters used for the experiment are shown in Table 1. It should be mentioned that with regard to the effective post curing time (less than 10 minutes) related to the printed parts in this experiment (low scale objects), the fixed post curing time of 5 minutes was applied for all parts.Table 1 MIP-SLA Process Parameters based on a Central Composite DesignKey process input variable Negative alpha (-α)Layer Thickness (µm)30UV sensitivity0.320Low450.370Center650.435High Positive alpha (+α)851000.5000.5509.5 33.1 8 7.62 9.53412.725.463.55Fig. 4 ASTM D638 Type 5 samples used for tensile test (Units: mm).17644.2 Shot peeningShot peening is a surface treatment process widely used to improve fatigue strength of metal parts by inducing compressive residual stresses in surface layers.

In order to have control of shot peening in metals, industry depends on the use of an “Almen gauge” introduced by J.O. Almen. Rectangular steel strips of controlled composition and thermal history are peened while being held flat. Deflection on release from the sample holder is measured. This deviation is the “Almen arc height” and is indicative of the amount of residual stresses in the subsurface. This arc height is used to plot a saturation curve. The saturation curve is plotted with the Almen strip deflection as a function of peening time. This curve indicates the point at which deflection in the material does not change by more than 10%, even if the peening time/peening pressure is increased. Since Almen strips pertain strictly to metals, the term pseudo-Almen strip was used to describe MIP-SLA printed polymer strips used to determine saturation. The geometry of the pseudo-Almen strip is presented in Fig.

The geometry of the pseudo-Almen strip is presented in Fig. 5 and is geometrically based on a standard Almen strip from Electronics Inc.76.09818.987Pseudo-Almen test strip1.3Fig. 5 Pseudo-Almen strip geometry used to determine saturation by shot peening an MIP-SLA polymer (all dimensions are in mm).There exist a standard procedure for finding the saturation point for metals. In the case of polymers, there is no standard defined for shot peening. Without a standard procedure, a methodology similar to metals was implemented to find the saturation curve for polymers. Pseudo-Almen strips were printed on the B9 creator and shot peened. The deflection of the pseudo-Almen strips was measured using images taken with a Nikon D3200 DSLR camera and analyzed using image processing techniques in AutoCad software.

The deflection of the pseudo-Almen strips was measured using images taken with a Nikon D3200 DSLR camera and analyzed using image processing techniques in AutoCad software. Pseudo-Almen strips were shot peened with glass beads from 10 to 25 passes in increments of 5 passes at a constant maximum air pressure of 85 psi with the gun nozzle one inch from surface.At these shot peening parameters, there was no visual deflection observed in the pseudo-Almen strips. Since deflection is required to plot a saturation curve, the shot peening media was changed to steel since steel beads provide greater impact energy to impart compressive stresses. Coverage was also increased to 60 passes using the same shot peening parameters described previously. Even with increased coverage and steel beads, there was no visual deflection in the pseudo-Almen strips (Fig. 6).

Even with increased coverage and steel beads, there was no visual deflection in the pseudo-Almen strips (Fig. 6). In order to determine if the material had any ability to plastically deflect, a strip was shot peened for a set time of 5 minutes with steel beads at maximum pressure with the nozzle directly on the material’s surface. Still, there was no deflection observed in workpiece. The material was too flexible. The hyper elastic nature of the material prevented permanent plastic deformation.1765Pre-shot peeingPost-shot peeing10 mmFig. 6 Pseudo-Almen strips before and after shot peening. No deflection was observed.As there was no deflection observed under any shot peening parameters for this photopolymer, a saturation curve obtained for ABS material was used.

No deflection was observed.As there was no deflection observed under any shot peening parameters for this photopolymer, a saturation curve obtained for ABS material was used. Using this saturation curve, the ASTM samples were shot peened on both sides for 20 passes using steel beads with one pass defined as full coverage; the shot peening nozzle was kept 0.5 inch from the sample.4.3 Tensile TestingUltimate strengths of shot peened ASTM D638 Type 5 samples were measured using an MTS 810 tensile tester (see Fig. 1). The grips of tensile tester were tightened and aligned between the two grips of unit. A computerized load cell located in the grips of the tensile tester measured applied force to the gauge section of the samples. To calculate strain, the distance between the grips was measured for each sample, and elongation was obtained from computer software. Engineering stress was calculated by dividing the force applied on the sample with the cross-sectional area.

Engineering stress was calculated by dividing the force applied on the sample with the cross-sectional area. The cross-section for each sample was measured using digital vernier calipers. Each measurement was repeated three times and the average value was reported.5. Results5.1 Effect of Process Parameters on Build QualityAn analysis of variance (ANOVA) was performed on the tensile strength of the ASTM D638 Type 5 samples. Fig. 7 represents the interaction plot for ultimate tensile strength (UTS). Based on the interaction plot, the best print settings that provided the highest UTS were using a 45-micron layer height along with a UV sensitivity factor of 0.370. It can be concluded that by increasing the UV sensitivity, the UTS generally decreased.As per the ANOVA results in Table 2, layer thickness was the only parameter that had a statically significant effect on cross-sectional area with a P-value less than 0.05.

The P-value for both layer thickness and UV sensitivity were greater than 0.05 indicating the two parameters did not have any combined effect on UTS of polymer parts. That is, no factor significantly affects the UTS with 5% significance level.1766Layer thickness) a P M353345 µm 65 µm 85 µm(S T U31290.30.4 0.5 UV sensitivity0.6Fig.

The surface was shot peened according to the procedure identified in Section 4.2. From the design of experiments mentioned above, eleven different ASTM D638 samples were printed on the B9 creator using Monocure blue resin. Each sample was repeated three times and the average results were reported.Fig. 8a shows a comparison of the UTS for shot peened and non-shot peened samples with respect to layer height. From this figure, it can be observed that for most of the samples (except at a layer height of 30 µm), shot peening reduced the UTS as compared to the non-shot peened case. The dashed lines indicate the grand means. The difference between shot peened and non-shot peened samples was dependent on layer thickness. At lower layer thicknesses, this difference was greater compared to the 85 µm and 100 µm layer thicknesses.Similarly, the variation of ultimate tensile strength with UV sensitivity is shown in Fig. 8b.

8b. Here also, shot peened samples exhibited a decrease in tensile strength for most of the samples. Generally, the difference between shot peened and non-shot peened samples decreased as the UV sensitivity increased. The exception where shot peening improved the UTS was at a UV sensitivity of 0.5.1767P0.218 0.218 0.785 0.0933334Non-Shot peened Shot peened3233) a P M31) a P M32((S T U30S T U3129Non-Shot peened Shot peened302829(a)2070 Layer height (µm)1200.3(b)0.4 0.5 UV sensitivity0.6Fig. 8 Ultimate tensile strength vs. (a) layer height and (b) UV sensitivity for shot peened and non-shot peened MIP-SLA printed samples.With respect to both layer height and UV sensitivity, the average tensile strengths of samples decreased compared to non-shot peened samples. This may be due to damage introduced to the samples by shot peening.

This may be due to damage introduced to the samples by shot peening. Beads travelling at high velocities may have caused damage in the form of micro-cracks on the surface of the material rather than inducing favorable compressive residual stresses. Furthermore, it was interesting that some printing parameters were able to see improvements to strength after shot peening. The mixed results indicate that further investigation is needed on multiple polymer systems in order to draw wider conclusions related to the effect of shot peening on 3D printed polymers.So to find the effect of shot peening on polymer parts, ANOVA was performed on all printing conditions of both shot peened and non-shot peened samples. Table 3 below represents the ANOVA test on these two groups. Per the large P-value (0.074), which is greater than the accepted rate of 0.05, application of shot peening did not significantly improve or deteriorate the UTS of MIP-SLA parts.

Per the large P-value (0.074), which is greater than the accepted rate of 0.05, application of shot peening did not significantly improve or deteriorate the UTS of MIP-SLA parts. More data or a wider confidence level is needed to draw solid conclusions regarding the effects of shot peening on additive polymer parts.Table 3 One-way ANOVA Comparing UTS of Shot Peened and Non-shot Peened SamplesSource DF SS Factor Error TotalMS 6.22 3.29 0.074FP1 60 113.28 1.89 61 119.56.226. Summary & ConclusionsShot peening of pseudo-Almen strips made of a photopolymer using mask image projection stereolithography (MIP-SLA) did not cause any measurable deflection. This may be because shot peening did not induce any residual stresses because the work piece to hyper elastic. Instead, the1768saturation point of an alternative polymer (ABS) from a parallel project was used to determine the preferred shot peening condition on Monocure blue resin ASTM D638 Type 5 tensile test samples.

ANOVA analysis was done to find effect of layer thickness and UV sensitivity. From the analysis, layer thickness affected the cross section of the material. At a lower layer thickness, the layer bonded to previous layers more effectively thus improving dimensional accuracy of parts. ANOVA analysis was also performed on tensile test data of non-shot peened samples. The P-value was not less than 0.05 for both process parameters, which indicated that layer height and UV sensitivity did not have a statistically significant effect on tensile strength. This is in contradiction to what was found literature. This contradiction may due to the fewer number of replications used for the experiments that will affect the accuracy of the statistical analysis.Tensile tests were performed on shot peened samples, and results indicated that the average ultimate tensile strength (UTS) of shot peened samples was less than the non-shot peened samples.

An ANOVA analysis compared all the non-shot peened samples with all shot peened samples. From the analysis, it was found that the P-value was not less than 0.05 implying that shot peening did not have a statistically significant effect on the UTS. Based on the results, the effect of shot peening on the strength is inconclusive, but generally it will decrease. More studies using a wider variety of materials and printers as well as an investigation in to damage mechanisms on polymers from shot peening is needed.7. AcknowledgmentsUndergraduate and graduate students conducted this research as a course project in MECH 498/898: Additive Manufacturing at the University of Nebraska-Lincoln (UNL). The authors would like to acknowledge the support of the Department Chair in Mechanical and Materials Engineering at UNL, Dr. Jeffrey Shield, for enabling Drs.

The authors would like to acknowledge the support of the Department Chair in Mechanical and Materials Engineering at UNL, Dr. Jeffrey Shield, for enabling Drs. Rao and Sealy to develop a research-based course in additive manufacturing in order to spur more undergraduate involvement in research.8. References[1] Saleh, N., Hopkinson, N., R.J.M. Hague, Wise, S., 2004. Effects of electroplating on the mechanical properties of stereolithography and laser sintered parts, Rapid Prototyping Journal, 105, pp. 305-315.[2] Zarringhalam, H., Hopkinson, N., 2003. Post-Processing of Duroform Parts for Rapid Manufacture, Solid free form fabrication symposium, University of Texas-Austin, pp. 596- 606.[3] Fuentes, F., Gallardo Jr, A., Martinez, J., Bulman, J., Foyos, J., Mendelson, M., Noorani, R., Fritz, B., 2002. Effects of Cryogenic Aging on a Rapid Protoyped polymer, International Solid Freeform Fabrication Symposium, Austin, Texas, pp.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of Process Parameters on Selective Laser Melting Al2O3-AlCermet MaterialHailong Liao*, Haihong Zhu*, Junjie Zhu†, Shijie Chang†, Xiaoyan Zeng* *Wuhan National Laboratory for Optoelectronics, Huazhong University of Science and Technology,Wuhan, Hubei 430074, P. R. of China†Shanghai Aerospace System Engineering Institute, Shanghai 201108, P.R. of ChinaAbstractThe cermet composite material is one of the researches focuses in the field of materials, for it can combine the toughness of metal and the hardness of ceramics. In this work, Al2O3-Al cermet composite with a mass ratio of 1:1 was fabricated by selective laser melting process. The effect of process parameters on the relative density and Al2O3 loss rate, as well as the Al2O3 loss mechanism, was investigated in detail.

The effect of process parameters on the relative density and Al2O3 loss rate, as well as the Al2O3 loss mechanism, was investigated in detail. The results show that Al2O3 undergoes melt recrystallization and is significantly aggregated. The aggregated Al2O3 exhibits a network distribution in the metal matrix. The process parameters have a great influence on the relative density and the Al2O3 loss rate. As the scanning speed decreases, the relative density and the Al2O3 loss rate are changed with a contrary tendency. The loss mechanism is that the aluminum acts as a reducing agent, causing the Al2O3 becoming a gaseous substance Al2O during selective laser melting process.IntroductionAl2O3 ceramics have excellent wear resistance and corrosion resistance, which are rich in resources and low in price[1,2]. Therefore, the development and application of Al2O3 ceramic materials have a great economic effect.

Therefore, the development and application of Al2O3 ceramic materials have a great economic effect. However, the brittleness of ceramics seriously restricts the development of their application[3]. Recently, the development of cermet material has made up for the lack of plasticity of Al2O3 ceramics. The cermet is composed of metal and ceramic, and the latter has a volume fraction of 15 to 85%[3]. Aluminum is one of the most widely used metals with high strength to weight ratio and low densities. It is also commonly used in the fabrication of cermet[4]. Al2O3-Al cermet materials have both excellent properties, such as high specific strength, high specific modulus, good wear resistance, etc[5,6]. They are promising in industry. Conventional methods for fabricating Al2O3-Al cermet materials generally include stir casting[7,8], hot isostatic pressing[9,10], and sintering[11,12].

They are promising in industry. Conventional methods for fabricating Al2O3-Al cermet materials generally include stir casting[7,8], hot isostatic pressing[9,10], and sintering[11,12]. However, these methods limit the application of Al2O3-Al cermet due to their inability to form part with complex structure. It is urgent to develop a new method for fabricating cermet part with complex structure.Selective laser melting (SLM) is now widely used in the aerospace industry because of its ability to form complex shaped parts[13–15]. Studies have shown that this method can also be applied to the manufacture of cermet materials[16,17]. This paper investigates the effect of process parameters on the selective laser melting of Al2O3-Al cermet with a mass fraction of 50wt%.Materials and experiment proceduresMaterials and SLM processThe polygonal Al2O3 powder used in this experiment has an average particle size of 26.6 (cid:541)m.

The aluminium alloy powder is the AlSi10Mg with an average particle size of 33.1(cid:541)m. Ball milling was performed in a tumbling ball mill to obtain uniform Al2O3-AlSi10Mg composite powder, according to the weight ratio of9601:1(Fig.1). It reveals that the morphology and dispersion state of the mixed Al2O3-AlSi10Mg powders do not change markedly under the condition of mechanical blending.Fig.1. SEM image of the mixed Al2O3-AlSi10Mg composite powders after ball milling The experiments were carried out on a self-developed SLM-100 machine that equipped with an IPG YLR-200 fiber laser ((cid:540)=1.07(cid:541)m) with a maximum output power of 200W, an automatic powder layering appliance and an inert argon gas protection system. The SLM machine was described in detail in our previous publication[18]. The specimens were to be prepared on an AA 2024 aluminum alloy substrate with dimensions of 140mm×140mm×35mm.

The SLM machine was described in detail in our previous publication[18]. The specimens were to be prepared on an AA 2024 aluminum alloy substrate with dimensions of 140mm×140mm×35mm. The process was carried out under argon atmosphere with the concentration of O2 controlled below 200ppm.The sample with a size of 5mm×5mm×5mm was preliminarily fabricated using several parameters setsin order to study the effect of processing parameter. The used detailed parameters are shown in Table 1.Table 1 Processing parameters used in this study. Manufacturing parameters Laser power P, W Scanning speed v, m/s Layer thickness t, (cid:586)m Hatching space h, mmValue 200 150,200,250,300 and 350 20 0.05, 0.10CharacterizationThe relative density and Al2O3 loss rate of the samples were evaluated by an image process algorithm (IPA) based on C# language. This method was described in detail in our previous study[17]. The Al2O3 loss rate (S) is determined by Eq.

Phase identification was performed by X-ray diffraction (XRD) using a PANalytical X´pert PRO with Cu (cid:549)(cid:573) radiation.(cid:28595)On all averaged values, a confidence interval of 95%, based on the variance (cid:2026)(cid:3041)(cid:2879)(cid:2869), is used to indicate the spread.Results and discussionsMicrostructureFig.2 shows the SEM image of the vertical cross of SLMed sample fabricated under the speed of 350mm/s and hatching space of 0.10mm. The white substance is Al2O3. It is apparent that the Al2O3 undergoes melting and recrystallization. The molten Al2O3 aggregates in the molten pool and exhibits a network distribution. The red arrow identifies the unfused pore near the Al2O3 in Fig.2a. As an additive, Al2O3 affects the flow of AlSi10Mg in the molten pool, especially the molten Al2O3, which easily hinders the flow of the molten pool. Thereby, unfused pores are formed. As shown in Fig. 2b, the aggregated Al2O3 ceramics have microcracks.

Thereby, unfused pores are formed. As shown in Fig. 2b, the aggregated Al2O3 ceramics have microcracks. It can be seen that the ceramic still exhibits strong brittleness. Of course, this is also related to the SLM being a fast cooling rate and high-temperature gradient. Al2O3 is prone to form micro-cracks during the process.Fig.3 is a photomicrograph of SLMed sample under different process parameters. The process parameters have a great influence on the distribution state of Al2O3. When the hatching space is 0.05 mm and scanning speed is 150mm/s, Al2O3 exhibits local aggregation, as shown by the white area in Fig. 3a-1. As the scanning speed increases, the Al2O3 distribution is more uniform and the content of Al2O3 in the matrix increases. When the hatching space is 0.10mm, the Al2O3 exhibits a network distribution, especially when the scanning speed is greater than 250 mm/s.Fig.4 is the XRD diffraction pattern of the typical SLM formed Al2O3-AlSi10Mg composite.

The XRD results show that the diffraction peak of Al2O3 decreases with the increase of the energy density. This confirms the occurrence of Al2O3 loss. As the energy density increases, i.e. the hatching space or scanning speed decreases, the Al2O3 loss increases significantly, as shown in Fig.3a-1 and Fig.3b-5. A uniform distribution of Al2O3-AlSi10Mg composites is obtained when scanning speed is in the range of 250 mm/s to 350 mm/s and the hatching space is 0.10mm.Fig.2. SEM image of the vertical cross of SLMed sample fabricated under the scanning speed of 350mm/s and hatching space of 0.10mm962Fig.3. Polished cross-sections of SLM-fabricated Al2O3-AlSi10Mg sample under varying process parametersAlSi + AbO3,,--..Zl ·a ,£ ro '--' .c ·;;; s::: V s::: ......~l50mm/s 0.05mm *+++ * + ...++• + +350mm/s 0.10mm2030405060708090Diffraction angle 20 (degrees)Fig.4.

The XRD diffraction pattern of typical SLM formed Al2O3-AlSi10Mg compositeRelative density and Al2O3 lossFig.5 shows the relative density of SLMed samples at different scanning speeds. As the scanning speed increases, the relative density slowly increases. When the hatching space is 0.10mm and the scanning speed is 250mm/s, the highest relative density can be achieved to 94%. When the scanning speed is greater than 250mm/s, higher than 93% relative density also can be achieved. This is because at low scanning speed, the energy density is high and a large number of pores appear. The pores are formed at low scanning speeds from gases trapped within the molten pool or evolved from the powder during solidification[19]. The distribution of Al2O3 is also more uniform as the scanning speed increases as shown in Fig.3, which reduces the possibility of defects. Noteworthy that as the scanning speed continues to increase, the relative density may decrease.963I 00.

Apparently that the relative density and the Al2O3 loss rate show a contrary tendency. As the scanning speed and hatching space increase, the loss rate of Al2O3 gradually decreases. When the hatching space is 0.10mm and the scanning speed is higher than 250 mm/s, the Al2O3 loss rate is less than 7.6%. It has been confirmed that Al2O3 and aluminum undergo redox reactions under high temperature conditions. Al2O3 is easily reduced by aluminum to form Al2O gas and overflows the molten pool[17].100~ 0 .....,, E 03 ..... "' "' .Q 0 ~806040. - - -' '■200o.05mm - -+- -0.I0mm''150200300 Scanning speed (mm/s)250--·350Fig.6. The Al2O3 loss rate under varying scanning speed and hatching spaceConclusionsThe microstructure of SLMed Al2O3-AlSi10Mg cermet material with a mass ratio of 1:1 were investigated. The Al2O3 ceramic in Al2O3-AlSi10Mg undergoes melting recrystallization under laser irradiation.

The Al2O3 ceramic in Al2O3-AlSi10Mg undergoes melting recrystallization under laser irradiation. When the hatching space is 0.05 mm, the Al2O3 tends to locally aggregate. As the scanning speed or hatching space increases, the Al2O3 gradually disperses and exhibits a network distribution. The effect of process parameters on the relative density and Al2O3 loss rate was also studied. The relative density and the Al2O3 loss show with a contrary tendency. As the hatching spacing or scanning speed increases, the density increases and the Al2O3 loss decreases. The Al2O3 loss is caused by the reduction of the aluminum alloy at high temperature to form Al2O gas.964Future worksAs that process parameters affect the loss and distribution of Al2O3. Future work will focus on how to avoid loss and agglomeration of Al2O3.

However, very few studies mention the vibration characteristics of FFF printed structures. This paper investigates the effect of four process parameters, raster angle, nozzle temperature, layer height and deposition speed, on the vibration properties of FFF printed Polylactic Acid (PLA) structure through modal analysis and design of experiment. The effects of all four parameters show a good agreement on the first fives modes of resonance. It was found that raster angle significantly affects both resonance frequency (16.6%) and loss factor (7.5%). Meanwhile, the impact of the other three parameters is relatively low (less than 4%), which is different from previous research results on static mechanical properties.

Meanwhile, the impact of the other three parameters is relatively low (less than 4%), which is different from previous research results on static mechanical properties. All these results provide a guidance for further application of FFF in vibration field.Keywords: Vibration properties, Fused Filament Fabrication (FFF), Modal Analysis, Polylactic acid (PLA), Design of Experiments (DoE)A substantially improved version of this paper appears in a special issue of the TMS publication JOM, March 2022.IntroductionFused Filament Fabrication (FFF), also known as Fused Deposition Modeling (FDM)which is a trademark, is one of the most widespread Additive Manufacturing (AM) technologies. Although it is initially known as a low-cost rapid prototyping technology, nowadays FFF has become much more powerful due to various technical advancements in the last decades[1].

Although it is initially known as a low-cost rapid prototyping technology, nowadays FFF has become much more powerful due to various technical advancements in the last decades[1]. Plenty new materials have been developed for FFF such as metallic [2], ceramic[3], fiber-reinforced polymers[4], and bio-polymers [5]. Recent developments in hybrid FFF [6] and multi-material FFF [7] provide a much wider space for the design and manufacturing of multi-functio nal structures. At the structure level, the design novel metamaterials and architecture structures [8], [9] with customized material properties has been realized via FFF.All these progresses enable the application of FFF in many fields, such as vibration control.Some preliminary attempts have already been done in the literature. Compton et al. [10] reported the extrusion-based 3D printing of cellular composites by a new epoxy-based ink with controlled521alignment of multi-scale, high aspect ratio fiber reinforcement.

Compton et al. [10] reported the extrusion-based 3D printing of cellular composites by a new epoxy-based ink with controlled521alignment of multi-scale, high aspect ratio fiber reinforcement. Matlack et al. [8] demonstrated the fabrication of a metamaterial for low-frequency and broadband vibration absorption with FFF printed lattice and manually inserted steel cube damper. Mizukami et al. [4] presented a FFF printed locally resonant carbon-fiber reinforced acoustic metamaterial for attenuation of broadband vibration.Despite the research efforts, the application of FFF in the field of vibration control is still very few compared with its application in the other fields like static mechanical. As reported, the FFF process is controlled by many process parameters which have a large impact on the final parts’ mechanical properties [11]–[14].

As reported, the FFF process is controlled by many process parameters which have a large impact on the final parts’ mechanical properties [11]–[14]. Thus, one of the biggest obstacles remains in the lack of research on the effect of process parameters on the vibration properties of FFF printed structures.In the literature, large efforts have been made to discover the relationship between the process parameters and static mechanical properties of FFF printed parts, including tensile strength [11], flexural strength [13], fracture properties [14] and fatigue [15]. Rezayat et al. [16] reported that a raster angle along the direction of load resulted in higher tensile strength because the bonding of polymer between the adjacent filament is not as strong as bonding of polymer within the filament. Goh et al.

Goh et al. [12] demonstrated that lager layer height and higher printing speed result in smaller degree of cooling and higher average temperature of the deposited filament, which can improve the interlayer fusion and results in higher mechanical properties in across-the-layer direction.Concerning the vibration properties, although many studies mentioned the vibration characterization of FFF printed structures [14] [18] [19], very few articles aimed at finding the relation between the process parameters and the vibration properties of final parts. Seedhara et al. [20] investigated the effect of build orientation and layer height on the modal properties of FFF printed ABS beams. They reported that the increase of layer height declined the natural frequencies and the specimen printed in horizontal orientation had the highest natural frequencies. Raffic et al.

They reported that the increase of layer height declined the natural frequencies and the specimen printed in horizontal orientation had the highest natural frequencies. Raffic et al. [21] studied the effect of infill density, layer height and deposition speed on the natural frequency and resonant amplitude of FFF processed Polyethylene Terephthalate Glycol-modified (PET-G) and Acronitrile Butadiene Styrene (ABS) beams.A comprehensive understanding of the effects of process parameters on the vibration properties of FFF fabricated parts is essential for further applications of FFF in vibration control. The purpose of this paper is to evaluate the performance of FFF in terms of vibration properties and characterize the effect of four process parameters, raster angle, nozzle temperature, layer height and deposition speed, on the vibration properties of FFF printed structures.

Firstly, the reproducibility of the proposed protocol is evaluated in terms of the measured vibration properties of printed structures. Secondly, the effect of four key process parameters on the vibration properties of FFF printed PLA beams is investigated, where Design of Experience (DoE) and modal analysis are used. Finally, all these results provide a guidance for further application of FFF in vibration field.522Protocol of InvestigationIn this section, the protocol of the investigation of the process parameters on thevibration properties of FFF printed PLA beams is presented.Phase I: Set up of FFF and DoEMaterial, machine and geometryThe material used is Polylactic Acid (PLA) (Raise Premium 1.75 mm PLA filament), which is the most used materials in FFF well known for its high rigidity, high strength, low melting point, non-toxic, non-irritation, and good biocompatibility [22].

PLA’s high stiffness, good adaptability and operation simplicity enlarge its potential application in vibration control. The 3D printer used is a commercial dual extruder FDM machine Raise 3D E2. To investigate the flexural vibration modes, a beam specimen with a geometry of 3 mm thickness, 20 mm width and 200 mm length is chosen.Parameters and their variation ranges for DoEFour process parameters and one of their interactions are selected here as factors in design of experiment: raster angle (R); nozzle temperature (T); layer height (H); deposition speed (V); interaction between raster angle and nozzle temperature (R-T). In the literature, raster angle has a severe effect on the mechanical resistance of printed structure because of the anisotropy nature of FFF process [11]. The nozzle temperature, layer height and print speed control the thermal history of printing which dominants the bonding strength between filaments [12].

The nozzle temperature, layer height and print speed control the thermal history of printing which dominants the bonding strength between filaments [12]. According to previous research, PLA has a high anisotropic property induced by the nature of the deposition process which makes the bond strength between filaments much weaker than the strength of deposed filaments [23]. The raster angle can change the orientation of structure anisotropy by arranging the direction of adjacent filament bonding. Meanwhile, the nozzle temperature influences the thermal history of printing and eventually control the filament bonding strength. Therefore, it’s supposed that there is an interaction between the raster angle and the nozzle temperature on the vibration properties of FFF structures.The limit values of parameters were determined through a preliminary test as shown in Table 1. Nozzle temperature, layer height and deposition speed have three levels.

Nozzle temperature, layer height and deposition speed have three levels. To study the effect of interaction between the raster angle and the nozzle temperature, a Taguchi orthogonal array was developed in the next section. Thus, the raster angle only has two levels due to statistical limitation of DOE method. To obtain the maximum difference 1 L18(2 in stiffness and verify the maximum effect of parameters, 90° and 0° were chosen as raster angles.

To obtain the maximum difference 1 L18(2 in stiffness and verify the maximum effect of parameters, 90° and 0° were chosen as raster angles. Values of other process parameters are presented in Table 2.3 3)Table 1 Levels of factors for Taguchi design of experiment of PLALevels R (°) T (°C) H (mm) V (mm/S)1 2 30 90 /190 220 2500.1 0.2 0.320 55 90523Table 2 Fixed Process parameters Parameter Build orientation Horizontal 0.4mm Extrude width Flowrate 100% Bed temperature 60°C 100% Cooling 0 Shell 100% Infill density Line Infill patternValueDesign of experiment in use of Taguchi methodTaguchi method is selected for the design of experiment because of its good balance between statistical analysis and operation simplicity [24]. A Taguchi orthogonal array is selected to evaluate the effect of these 5 factors.

A Taguchi orthogonal array is selected to evaluate the effect of these 5 factors. R has two levels, while T, H and V have three factor levels.1 L18(23 3)Table 3 TaguchiOrthogonal Arrays for the DoE of PLA3 1 Trials Specimen Replication 3 L18(2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2PLA 1111 1 PLA 1122 2 PLA 1133 3 PLA 1211 4 PLA 1222 5 PLA 1233 6 PLA 1312 7 PLA 1323 8 9 PLA 1331 10 PLA 2113 11 PLA 2121 12 PLA 2132 13 PLA 2212 14 PLA 2223 15 PLA 2231 16 PLA 2313 17 PLA 2321 18 PLA 2332)R 1 1 1 1 1 1 1 1 1 2 2 2 2 2 2 2 2 2T 1 1 1 2 2 2 3 3 3 1 1 1 2 2 2 3 3 3Factors H 1 2 3 1 2 3 1 2 3 1 2 3 1 2 3 1 2 3V 1 2 3 1 2 3 2 3 1 3 1 2 2 3 1 3 1 2R-T(1,1)(1,2)(1,3)(2,1)(2,2)(2,3)Phase II: experimentVibration test setup: modal analysis524Modal analysis is used to characterize the vibration properties of PLA beams.

Figure 1 shows the experimental setup for measuring the frequency response curves of beams in forced vibration.The shaker generates vibrations that were controlled and measured by a controller device (UCON system). The range of frequency excitation is between 10 and 2000 Hz. The vibration amplitude of the shaker was controlled using an accelerometer (PCB Piezotronics). The amplitude of the shaker was chosen to minimize nonlinear vibrations typically 1 mm displacement between 10 and 2000 Hz and a 1g acceleration between 10 and 1500 Hz. The amplitude of the beam's vibration was measured by a laser vibrometer.

The amplitude of the beam's vibration was measured by a laser vibrometer. The response spectrum using FRF (Frequency Response Function) data was used to determine the resonant frequencies and the loss factors which was calculated by the Half Power Bandwidth Method (HPBM, also known as-3dB method) [25].Figure 1 Experiment setup of modal analysisFor PLA specimens, the boundary condition is clamped-free and the laser spot was pointed on the free extremity of the beam which can capture all vibration mode on the selected spectral. The beams were attached to the shaker by an apparatus with a clamp length of 15 mm.Figure 2 Boundary condition of modal analysis. Clamped-FreeMeasured vibration properties525To establish a quantitative qualification, two vibration properties were selected as key factors: the resonant frequency and the loss factors. Resonant frequency is a fundamental index of vibration in frequency domain vibration test [25].

Resonant frequency is a fundamental index of vibration in frequency domain vibration test [25]. Moreover, the damping of structures is very important for vibration control. And loss factor is a common index used to characterize the damping of material or structure, which is equal to the ratio between the loss modulus and the storage modulus [26].Result and DiscussionReproducibility evaluation of the proposed protocolTo guarantee the reliability of the identified parameters effect, the precision of the FFF process and the vibration test prototype was evaluated in terms of the reproducibility of the measured vibration properties. PLA beams with a geometry of 3 mm thick, 20 mm large and 200 mm long were printed six times repetitively, using medium parameter levels as follows: R, 0°, T, 220°C; H, 0.2 mm; V, 55 mm/s.As shown in table 4, percentage error is the ratio between the standard deviation and the average value.

The results of resonant frequencies show a high-level reproducibility with a percentage error smaller than 1%. On the other hand, the percentage errors of loss factors are much higher than those of resonant frequencies, but still good (smaller than 6%). That is because the damping of PLA is relatively low. Therefore, the value of loss factors is too small to be precisely measured.Table 4 Reproducibility of FDM printed PLA beam.Item Resonant Frequency (Hz) 21.97±0.12 Percentage error (%) Loss Factor Percentage error (%)Mode 1Mode 2 135.77±1.02 0.75Mode 3 379.52±2.45 0.65Mode 4 744.56±4.77 0.65Mode 5 1228.84±8.12 0.670.57 0.0292±3.65E-04 0.0149±8.18E-04 0.0125±6.91E-04 0.0122±6.53E-04 0.0127±6.31E-04 1.265.865.785.565.13With low level deviation and percentage error, the reproducibility of fabrication and modal analysis is verified within the selected process parameter level.

The reproducibility may vary with the changes of process parameter, but this variation is the influence of process parameter itself but not the effect of machine calibration nor the effect of experiment setup.Main effects on the resonant frequency of PLA beamsThrough modal analysis, five resonant modes of PLA beam are obtained. The mains effects of parameters on PLA are presented in Figure 3, 4 and 5. Since the resonant frequency of each mode varies from 20 to 1500 Hz, the mains effects of parameters are normalized by the average resonant frequency of all specimens in each mode to make a comparable analysis. Similarly, main effects of loss factors are normalized by the average loss factor of all specimens in each mode.Although the effects of the other three parameters are relatively low (below 3%) compared with the effect of raster angle (about 16%).

These effects show a consistent trend526across all five resonant modes.Considering raster angle in Figure 3 (a), from 90° to 0°, the resonant frequencies decrease by 16.6%. This shows that PLA has high anisotropic property in terms of resonant frequency, which is consistent to the reported anisotropy [27], [28] in static mechanical properties.

The increase of resonant frequency from low temperature to middle temperature can be explained by the improve of bonding degree between filaments. As reported by Goh et al. [12], a higher average temperature leads to better bonding strength. The reason of the decrease527of resonant frequency from middle temperature to high temperature is believed to be radical degradation [22], where the high temperature level (250°C) is 80°C higher than the melting temperature of PLA [29].Resonant frequency increases with the increase of layer height as shown in Figure 3 (c). From 0.1mm to 0.2 mm, resonant frequency increases by 2.2%; from 0.2 mm to 0.3 mm, resonant frequency increases by 0.7%.

From 0.1mm to 0.2 mm, resonant frequency increases by 2.2%; from 0.2 mm to 0.3 mm, resonant frequency increases by 0.7%. As layer height increases, fewer number of layers need to be deposed, resulting in fewer thermal cycles and smaller thermal gradient during printing, which was reported to have a positive effect on adhesion strength between filaments [12], thus leads to higher stiffness and higher resonant frequency.For deposition speed, as shown in Figure 3 (d), the resonant frequency remains in the same level (0.7% of variation) as the deposition speed increase from 20 mm/s to 55 mm/s, and decreases by 1.5% as deposition speed increase from 55 mm/s to 90 mm/s. Higher deposition speed leads to fewer time for bonding formation between filaments and smaller necks.The effect of interaction between raster angle and layer height is shown in Figure 4.

Higher deposition speed leads to fewer time for bonding formation between filaments and smaller necks.The effect of interaction between raster angle and layer height is shown in Figure 4. The effect of nozzle temperature shows same trends on all levels of layer height but is more important at higher level of raster angle. At 0°, resonant frequency increases by 0.8% as nozzle temperature increases from 190 °C to 220°C then decreases by 0.4% as nozzle temperature increases from 220 °C to 250°C. This effect is very weak which is at the same level of measurement deviation assessed in reproducibility test. At 90°, resonant frequency increases by 2.3% as nozzle temperature increases from 190 °C to 220°C then decreases by 2.2% as nozzle temperature increases from 220 °C to 250°C, which is more significant than effect of nozzle temperature at 0°. That’s because the vibration of beam is in flexural mode.

That’s because the vibration of beam is in flexural mode. Thus, at 90°, the direction of filament bond is parallel to the axe of flexural mode and contributes more to the global flexural modulus of beam. This interaction shows that filament bond strength plays a more important role in vibration when the raster angle is parallel to the axe of vibration mode.For the resonant frequencies of PLA beam, the following insights were retrieved:Raster angle has the most significant effect on the resonant frequency of PLA beam;Effects of nozzle temperature, layer height and deposition speed are relatively low;Effects of parameters show a consistency impact on all five resonant modes.Main effects on the loss factors of PLA beamsSince the variation of loss factor for each mode is very similar, the following quantitative descriptions will be based on the first resonant mode. As shown in Figure 5 (a), raster angle has the most significant influence on the loss factor of PLA beams.

As shown in Figure 5 (a), raster angle has the most significant influence on the loss factor of PLA beams. From 0° to 90°, loss factors reduce by 7.5%. Considering nozzle temperature, loss factor drops by 0.7% from 190°C to 220°C and528increases by 1.4% from 220°C to 250°C.

For deposition speed, loss factor increases by 0.8% from 20mm/s to 50mm/s, then increases by 1.4% from 55mm/s to 90mm/s.r o t c a Fs s o L d e z i l a m r o N(a) Raster Angle(b) Nozzle Temperature(c) Deposition Speed(d) Layer HeightFigure 5 Effect of process parameters on the loss factors of PLA beamFor the loss factors of PLA beam, the following insights are retrieved:Values of loss factors are very small which indicates that the damping capacity of PLA beams is very low;Raster angle has a significant effect on the resonant frequency of PLA beam, while effect of nozzle temperature, layer height, raster angle is relatively weak;Except layer height, parameters shown consistency impact on all five resonant modes;Except for the layer height, effect of each parameter on the loss factors shows an inverse trend to its effect on the resonant frequencies.Analysis of VarianceIn order to provide a correct interpretation of the effect of factors, an ANOVA analysis is necessary.

codes : 0 <'***'< 0.001< '**' <0.01 <'*' <0.05Considering the effect on the loss factor (Table 6), the raster angle is the only significant parameter, but it should be noted that there are large residual errors (contribution: 40.17%). There can be two sources for this residual error: (1) variable process parameters not considered in fabrication; (2) noisy of measurement in modal analysis. If the main cause is the first one, the residual errors for the resonant frequency should have a similar value compared to that of loss factor, which is not the fact. Therefore, this unusual residual error is mainly caused by the noisy measurement of loss factors in modal analysis because small value of loss factors makes the noise of measurement much more important.Table 6 Significance of parameters on the loss factor of PLA beamSource Sum Sq. R 4.20E-05 1 T 9.44E-07 2 H 7.45E-07 2 V 2.49E-06 2 2.29E-06 2 R*T Error 2.06E-05 26 Total 6.91E-05 35d.f. Mean Sq.

R 4.20E-05 1 T 9.44E-07 2 H 7.45E-07 2 V 2.49E-06 2 2.29E-06 2 R*T Error 2.06E-05 26 Total 6.91E-05 35d.f. Mean Sq. F 4.20E-05 4.72E-07 3.72E-07 1.24E-06 1.14E-06 7.93E-0752.97 0.59 0.47 1.57 1.44Prob>F 9.93E-08 *** 0.56 0.63 0.23 0.25Signif Contribution (%)59.66 -0.93 -1.22 1.30 1.01 40.17 100.00Signif. codes : 0 <'***'< 0.001< '**' <0.01 <'*' <0.05Conclusion and PerspectiveIn this study, the effect of four process parameters, raster angle, nozzle temperature, layer height and deposition speed, on the vibration properties of FFF printed PLA beams are investigated in use of modal analysis and DoE. It’s found that FFF printed PLA beams have high resonant frequencies and small loss factors, which reflects its high stiffness and weak damping. A high reproducibility (percentage errors smaller than 1%) is also obtained in terms of resonant frequency in selected parameter level.

A high reproducibility (percentage errors smaller than 1%) is also obtained in terms of resonant frequency in selected parameter level. Therefore, PLA is more suitable to be used as a rigid material to reinforce the stiffness of structures to prevent vibration other than a damping material to dissipate vibration energy. Follow guidelines are presented for the future design and manufacturing to530optimize the resonant frequency of PLA based FFF printed structures:1. Raster angle has the largest effect on both resonant frequency (16.6% of variation) and loss factors (7.5% of variation). A raster angle perpendicular to the axe of vibration form (axial axe for flexural vibration mode) leads to a higher stiffness as well as higher resonant frequency.2. The effect of layer height on the resonant frequency is much smaller than that of raster angle but still significant.

The effect of layer height on the resonant frequency is much smaller than that of raster angle but still significant. Larger layer height leads to higher resonant frequency as the adhesion strength between filaments is improved by the reduction of thermal cycles and thermal gradient.3. The effects of nozzle temperature, deposition speed and the interaction between raster angle and nozzle temperature on the resonant frequency are also observed in the experiment but not statistically significant. Therefore, a constant manufacturing quality can be achieved with in the selected range of nozzle temperature and deposition speed in terms of vibration properties.It should be mentioned that the importance and influence of process parameters of different materials may be different. Therefore, future research should focus on materials different from PLA, such as elastomers, ceramics, and metals.

Therefore, future research should focus on materials different from PLA, such as elastomers, ceramics, and metals. The reproducibility of FFF process in terms of vibration properties should also be verified on different materials and geometries. Then a more universal guidance can be established to promote the further applications of FFF process in the manufacturing structures with complex geometry and multiple materials for vibration control.AcknowledgementsThe French Ministry of Research (DRRT), the regional Council “Région Lorraine”, the European Regional Development Fund (FEDER) has contributed to the funding of the vibration and wave propagation platform. Also, the authors would like to thank to the research platform Lorraine Fab Living Lab® for the technical assistance in this research.References[1] S. Singh, G. Singh, C. Prakash, and S. Ramakrishna, “Current status and futuredirections of fused filament fabrication,” J. Manuf. Process., vol. 55, pp. 288–306, Jul.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperThe effect of processing parameter on zirconium modified Al-Cu-Mg alloys fabricated by selective laser meltingXiaojia Nie*, Hu Zhang*,†, Haihong Zhu*, Zhiheng Hu*, Xiaoyan Zeng*Wuhan National Laboratory for Optoelectronics, Huazhong University of Science and Technology, Wuhan, Hubei, P.R. of China†School of Optical and Electronic Information, Huazhong University of Science andTechnology, Wuhan, Hubei, P.R. of ChinaAbstractThe newly designed alloy compositions for selective laser melting (SLM) have aroused great interest. In this study, zirconium modified Al-Cu-Mg alloys were fabricated by SLM. Results show that crack-free samples with relative density of nearly 100% were obtained by optimizing the processing parameters.

Results show that crack-free samples with relative density of nearly 100% were obtained by optimizing the processing parameters. With the increase of scanning speed, the relative density decreases due to insufficient energy input. In addition, the microstructure transforms from homogeneous to bio-modal, the reason is the unstable flows caused by the high scanning speed. The small hatching space will provide more energy input and preheat, leading to the coarse surface.IntroductionAdditive Manufacturing (AM) processes enable to fabricate parts with very complex shape[1]. Selective laser melting (SLM), one of the preferred AM techniques, has recently gained considerable attention due to high manufacturing flexibility, near- net-shape production and efficient use of raw material[2, 3]. Al-Cu-Mg alloy has become an attractive material due to the low density, high fracture toughness and fatigue strength[4, 5].

Al-Cu-Mg alloy has become an attractive material due to the low density, high fracture toughness and fatigue strength[4, 5]. Nowadays, processing high strength aluminum alloys by SLM is confronted with great difficulties because of its poor flowability, high reflectivity, high thermal conductivity, large solidification range and hot cracking susceptibility[6]. Therefore, the biggest stumbling block to the application of high strength aluminum in SLM is the hot crack. To expand the scope of aluminum alloys suiting for SLM, new high strength aluminum alloys are urgent to be investigated.Recently, scandium (Sc), zirconium (Zr), titanium (Ti) and vanadium (V) have been proved as the effective microalloying element for aluminum alloys to improve the weldability and mechanical properties[7-9]. Nowadays, the addition of microalloying element has been applied to SLM aluminum alloys to improve their formability.

Nowadays, the addition of microalloying element has been applied to SLM aluminum alloys to improve their formability. The microstructural evolution, densification, properties and heat treatment of Al-Mg-Sc-Zr fabricated by SLM have been investigated[10-13]. Zhang H et al.[14] have found that the Zr modified Al-Cu-Mg parts with ultrafine grain exhibits better mechanical properties and broader processing window. However, the effect of processing parameters on the Zr modified Al-Cu-Mg alloy has not been studied.1246The present work investigated the effect of scanning speed and hatching space on Zr modified Al-Cu-Mg alloy fabricated by SLM. The relative density, surface roughness, microstructure and element distribution were discussed under different processing parameters.Materials and experiment proceduresMaterialsThe spherical gas atomized Al-Cu-Mg powders with an average particle size of 36 μm was used in the experiments.

The powder had a composition of 93.23 Al, 4.24 Cu, 1.97 Mg, and 0.56 Mn in wt.%. The Al-Cu-Mg powder mixture with zirconium particles were blended by mechanical mixing in an argon atmosphere for 4 h. The morphology of pure Zr powders and the Al-Cu-Mg powders mixture with 2% zirconium are shown in Fig. 1.Fig. 1. Morphology of pure Zr powders (a) and Al-Cu-Mg powders with 2 wt.% Zr addition (b).SLM processingThe SLM experiments were conducted on a self-developed machine (LSNF-I) whose details have been given elsewhere[15, 16]. All samples were deposited on the commercially AA 2024 substrate in an argon environment with the concentrations of H2O and O2 controlled well below 20 ppm. The SLM processing parameters of cubic samples were given in Table 1.Table 1 SLM processing parameters used in the experiments.

The SLM processing parameters of cubic samples were given in Table 1.Table 1 SLM processing parameters used in the experiments. ExperimentParametersValueCubic samples Laser power (P, W) Dimensions (mm3) Scanning speed (V, m/min) Hatching space (HP, μm) Layer thickness (D, μm)200 5*5*5 5, 10, 15, 20, 25 and 30 60, 70, 80 and 90 40 90Scanning strategy (θ, °)1247CharacterizationThe particle size was measured using the Malvern UK Mastersizer 3000. The relative density of the cubic specimens was evaluated by image processing of eight cross-section optical micrographs using ImagePro Plus 6.0 software. All samples were subjected to a standard metallographic procedure. For microstructure analysis, samples were etched by a solvent which consists of 2.5 mL HNO3, 1.5 mL HCl, 1 mL HF and 95 mL deionized water. The microstructure was characterized using an optical microscope (OM, EPIPHOT 300).

The microstructure was characterized using an optical microscope (OM, EPIPHOT 300). The elements distribution of cubic samples was conducted by electron probe microanalyzer (EPMA). The surface roughness was obtained by using laser scanning confocal microscope (LSCM, KEYENCE VK- X200K).Results and discussionsFormabilityFig. 2 shows the effect of scanning speed on the relative density of the selective laser melted (SLMed) samples (hatching space fixed at 90 μm). It is observed that the relative density decreases with the increase of scanning speed. Nearly fully dense samples were obtained when the scanning speed is low enough (5 m/min).Fig. 2. Effect of scanning speed on the relative density.Fig. 3 demonstrates the cross-sections of samples fabricated at different scanning speeds (hatching space fixed at 90 μm). The irregular pores occur when the scanning speed reaches to 20 m/min.

3 demonstrates the cross-sections of samples fabricated at different scanning speeds (hatching space fixed at 90 μm). The irregular pores occur when the scanning speed reaches to 20 m/min. The relative density significantly decreases when energy input is insufficient, that is, the scanning speed is too high to provide enough energy to melt adequate metal powders. Therefore, the irregular pores occur due to the weak connection between the layers and tracks.1248Fig. 3. Polished cross-sections of samples fabricated at different scanning speeds: (a) 5 m/min, (b) 10 m/min, (c) 15 m/min, (d) 20 m/min, (e) 25 m/min and (f) 30 m/min.Fig. 4 shows the effect of hatching space on the relative density and surface roughness (scanning speed fixed at 10 m/min). From Fig. 4a-d, as the hatching space increases, the samples still fully dense. It clearly shows that the percentage of overlapping is adequate to guarantee the connection between the tracks.

From Fig. 4a-d, as the hatching space increases, the samples still fully dense. It clearly shows that the percentage of overlapping is adequate to guarantee the connection between the tracks. As far as surface roughness concerned, small hatching space (less than 70 μm) leads to high energy input and coarse surface with many metal balls, as shown in Fig. 4e and f. It is suggested that small hatching space enables the track fabricated previously to receive more energy and has preheating function to the following tracks, which generates the excessive size of the melt pool, therefore, balling phenomenon occurs. The surface roughness of samples fabricated at 60, 70, 80 and 90 μm are 10.47, 8.44, 6.97 and 6.46 μm, respectively. Taking into consideration the relative density, processing efficiency and surface roughness, the optimal processing parameters are laser power of 200 W, hatching space of 90 μm, layer thickness of 40 μm, and scanning speed of 10 m/min.Fig. 4.

4. Polished cross-sections and 3D images of samples fabricated at different hatching spaces (scanning speed fixed at 10 m/min): (a, e) 60 μm, (b, f) 70 μm, (c, g) 80 μm and (d, h) 90 μm.1249MicrostructureFig. 5 shows the microstructure of samples fabricated with different scanning speeds. The distribution of equiaxed grains is homogenous when the scanning speed is 5 m/min, as presented in Fig. 5a. Fig. 5b-c shows the typical bi-modal microstructure, that is, the fine equiaxed grains intersperse with coarse grains. However, the flow of molten metal is much more unstable when the scanning speed is above 15 m/min, as shown in Fig.5 d-f. More details can be found in the work of H. Zhang [14]. The boundary of melt pool becomes blurry due to the unstable flow. What’s more, the grain size slightly decreases with the increase of scanning speed. High scanning speed leads to the high cooling rate, which contributes to the refinement of grains.Fig. 5.

6d. It is worth noting that little zirconium oxides can be observed in the grains.Fig. 6. Quantitative chemical maps obtained using EPMA in the cross-section of samples fabricated at 10 m/min.ConclusionsEffect of scanning speed and hatching space on the relative density and surface roughness of SLMed 2 wt.% modified Al-Cu-Mg alloy was investigated. With the increase of the scanning speed, the relative density decreases. The hatching space influences the surface roughness of samples. The smaller hatching space, the coarser1250the surface roughness. The microstructure is affected by the scanning speed. The distribution of fine equiaxed grains transforms from homogenous to bi-modal due to the increase of the scanning speed. Al3Zr and ZrO particles serve as the nucleate during the solidification.Future worksAs that the suitable heat treatment is necessary for the SLMed samples is proposed.

Al3Zr and ZrO particles serve as the nucleate during the solidification.Future worksAs that the suitable heat treatment is necessary for the SLMed samples is proposed. The future works may concentrate on the influence of the heat treatment on the SLMed Zr modified Al-Cu-Mg alloy.AcknowledgementsThis research is supported by the National Natural Science Foundation of China(61475056) and the National Program on Key Basic Research Project of China (613281). The authors would like to thank the Analytical and Testing Center of HUST.References [1] A.B. Spierings, K. Dawson, P.J. Uggowitzer, K. Wegener, Influence of SLM scan- speed on microstructure, precipitation of Al 3 Sc particles and mechanical properties in Sc- and Zr-modified Al-Mg alloys, Materials & Design, 140 (2018) 134-143. [2] N. Kang, P. Coddet, Q. Liu, H. Liao, C. Coddet, In-situ TiB/near α Ti matrix composites manufactured by selective laser melting, Additive Manufacturing, 11 (2016) 1-6. [3] B. Vrancken, L.

 Processing Parameters In SLS Of Metal...Polymer PowdersB.Badrinarayan and J.W.Barlow Dept. of Chemical Engineering, The University of Texas at Austin, Austin, Texas- 78712,U.S.A.Abstractpaper describes the effect of processing parameters in SLS of metal-polymer powder mixtures. Test bars for measurement of strength and density were made from copper-PMMA powder mixtures using SLS. The effect of Energy Density during laser processing, effect of vector length, bed temperature, polymer melt index and initial binder content on part strength and density are discussed in this paper. The green part strengths and densities were found to increase with Energy Density upto a value of 6.0 caVcm2 and then they drop off due to polymer degradation. Parts made with lower vector lengths yielded higher strengths and densities than those with higher vector lengths.

Parts made with lower vector lengths yielded higher strengths and densities than those with higher vector lengths. Parts processed with a low melt index polymer binder (around 6.0 gmllOmin) showed higher strengths than parts processed with higher melt index binders. High strength values were obtained for green parts made from powders that had a greater initial binder content. Bed temperature did not have a very significant effect on part strength and densities.IntroductionSintering comes under the class of Solid Freeform Fabrication Processes for rapid Selective parts directly from a CAD model without part specific tooling or human manufacture intervention [1,2]. This technology has been developed at the University of Texas at Austin to produce polymer parts as well as indirect metal and ceramic parts that have polymer as an intermediate binder.

This technology has been developed at the University of Texas at Austin to produce polymer parts as well as indirect metal and ceramic parts that have polymer as an intermediate binder. There are a variety of material and processing issues involved in manufacture of green metal parts using an intermediate polymer binder, namely the polymer melt index, Energy density during laser processing of parts, vector length of part being fabricated, the bed temperature employed during processing and the content of polymer binder that is used for green part fabrication. This work is directed towards a better understanding of these various issues which developing a good process control system for the SLS process and ensure qualityhelp rePeatabilitythe manufacturing process.Experimental ProcedureCopper powder, -325 mesh (me.an size of 5-10Jlm) which was used along with a polymer binder in metal part fabrication by SLS, was obtained from Reade International, Rumson, New Jersey.

PMMA and copolymers of MMA and BMA were used as intermediate binders for the metal powder in SLS processing. The synthesis of these polymer emulsions has been discussed in by Vail [3]. The polymer was obtained in powder form by spray drying the emulsion polymer in an Anhydro Modell Spray Drying Unit. All SLS processing experiments were carried out by mixing the polymer powder with the copper powder in the desired proportion. SLS processing was carried out in a nitrogen environment to keep the processing chamber inert. A layer thickness of 0.005 inches was used in all the SLS processing eXPeriments. Three point bend specimens (3.0" x 1.0" x 0.25") were made by SLS for measurement of strength and density of green parts.55Effect OfPolymer important The Melt the laser strength of the parts ""'......"'.. """'~.... Index on strength separately spray processed by SLS. 20J.1sec, 16 to 56 grid strength of parts Increases high viscosity weight,AsScan UBJ,...."""JL...

BJ'''''''' .....agreement with the data of Vail [3] for copolymer coated A-5000 soda-lime glass powder (copolymers 80/20-4-S, 80/20-6-S, 80/20-8-S). Vail also found that when binder melt flows were substantially smaller (in the range of 0.07 to 0.96 gms/IOmin) the modulus of rupture for the green SLS bars were often lower than the green SLS bars made with binders that had higher Melt Indices. The effect of Melt Index on green density is shown in Fig 6. The parts made from all the three polymer binders have comparable densities ranging from 41 % to 46%. The higher the Melt Index the better the flow and therefore greater the sintering and the resulting green density.

The higher the Melt Index the better the flow and therefore greater the sintering and the resulting green density. However in SLS the effect of Melt Index on density is not very pronounced as for the most part the density seems to be determined by the interlayer porosity that occurs on layering powder.Effect Of Bed Temperature: In SLS processing the material directly under the laser beam is at a much higher temperature than the material that has already been exposed to the laser energy or the material unexposed to the laser beam. Also, the surface temperature of the powder under the laser is much higher than that of the material below the surface. The temperature gradients in the laser heated material can cause differential volume changes because the sintering rate is a function of temperature, and hence residual stresses can build up in the material.

The actual residual stress distribution in the sintered part depends on the temperature variations in the powder bed, the associated cooling rates, and the ability of the material to stress relax. Polymers show rapid stress relaxation at temperatures larger than their glass transition temperature.During SLS processing of powders the part side powder surface is usually heated to prevent curling of the part due to thermal stresses on laser irradiation. The metal-polymer powder systems the bed temperature is usually set close to the glass transition temperature of the polymer binder (about 2_3° below Tg). This helps in effective sintering and also removes curling. The effect of bed temperature on strength and density of green parts was studied for a copper copolymer mixed powder system. The copolymer composition was 40% by volume. Three different bed temperatures, 85,75 and 65°C, were used.

The copolymer composition was 40% by volume. Three different bed temperatures, 85,75 and 65°C, were used. The laser power was fixed at 9.5 Watts, the step period (SP) was set to 20Jlsec, the Scan Spacing (SCSP) was set at 0.005 inches, and the Step Size (SS) was varied from 16 to 57 grid units. Three point bend bars were made for strength tests. Figure 7 shows the effect of bed temperature on part strength. It is seen that the green part strength increases with the bed temperature, but this increase is limited to the glass transition temperature of the polymer. If the bed temperature is kept above the glass transition temperature of the polymer the entire powder bed cakes up and the SLS run would get aborted. Parts scanned at bed temperatures near the polymer Tg showed minimal or no curl compared to those scanned at lower temperatures. Figure 8 shows the % density of the green parts with bed temperature at different Energy Densities.

Figure 8 shows the % density of the green parts with bed temperature at different Energy Densities. We fmd that bed temperature has no significant effect on the green part density.Effect of Amount of Polymer Binder In SLS: Another factor that is to be considered in SLS processing of metals with a polymer binder is the amount of polymer to be used as binder. Lower amounts of polymer will be desirable if the aim is to go to fully dense metal parts with low shrinkages. If, however, the aim is to obtain a lightly sintered metal or oxidized metal preform with subsequent infiltration by another lower melting metal or a liquid epoxy resin then it is not necessary to ensure a low binder content in the green part. The effect of binder volume fraction on part strength is shown in Figure 9. This curve suggests that high binder content might be desirable for producing green parts with good strengths.

The effect of binder volume fraction on part strength is shown in Figure 9. This curve suggests that high binder content might be desirable for producing green parts with good strengths. As the amount of polymer increases there is greater sintering and bonding with the metal powder and hence there is an increase in strength for the green compact.57SummaryIn this paper the effects of material properties and processing issues have been explored. It is found that energy densities around 6.0 callcm2 yield parts with good green strengths. Green parts fabricated with low Melt Index PMMA showed greater strengths than those processed with higher Melt Index PMMA. Parts made with lower vector lengths yielded greater strengths than those made with higher vector lengths. However, it is important to note that too small a vector length could lead to polymer degradation and hence lower strength values.

However, it is important to note that too small a vector length could lead to polymer degradation and hence lower strength values. Bed Temperature does not have a significant impact on strength values in the temperature ranges considered in processing. However it is necessary to keep the bed temperature close to the glass transition temPerature (about 2_3 0 below Tg) to prevent warping of parts during SLS processing. Higher binder content in raw powder leads to greater strengths on processing. The results obtained in this work can be used as a guideline for SLS processing of metal parts with an intermediate polymer binder.AcknowledgmentThe authors gratefully acknowledge support for this work by DARPA / ONR Grant NOOO14-92-J 1394.References1. Carl Deckard and Joseph Beaman, 15th Conference on Production Research and Technology, University of California at Berkeley, p 623 - 630, (1989).2. Harris L. Marcus, Joseph Beaman, Joel W. Barlow and David L.

FECT OF PROCESSING PARAMETERS ON THE DENSITY AND MICROSTRUCTURE OF DIRECT LASER SINTERED Al-12Si POWDERS.E.O.Olakanmia, R.F.Cochranea, K.W.Dalgarnob. a Institute for Materials Research, University of Leeds, LS2 9JT. b School of Mechanical and Systems Engineering, Newcastle University, NE1 7RU, UK.AbstractThe effect of processing parameters on the sintering behaviour of gas atomised Al-12Si powders has been investigated. Laser power, scanning rate, scan spacing and layer thickness are found to control the densification and the resultant microstructural characteristics of the laser sintered parts. It was found that sintered density increased as the energy density increased reaching a maximum of 80.2% at an energy input per unit volume of 67 J mm-3.

It was found that sintered density increased as the energy density increased reaching a maximum of 80.2% at an energy input per unit volume of 67 J mm-3. For parts produced with a slightly lower power density (50 J mm-3), the microstructure consisted of fine dendrites with interconnected porosity while parts fabricated with a slightly higher power density (100 J mm-3) were noted to have a preponderance of coarse dendrites with a discontinuous network of irregular shaped pores surrounded by a fully dense aluminium-silicon matrix.Keywords: Al-12Si alloyed powder; Microstructure; Energy Density; Scanning Electron Microscopy (SEM); Densification.1. IntroductionThe specialty of the direct selective laser sintering (SLS) process has been identified as one of the most promising areas for the production of engineering components.

IntroductionThe specialty of the direct selective laser sintering (SLS) process has been identified as one of the most promising areas for the production of engineering components. This is because it does not require moulds to produce intricate geometries and eliminates the post-processing steps thereby resulting in the shortening of the production cycle as well as the reduction of production costs. For example, studies on the direct SLS of high speed steels (HSS) have been carried out [1-7]; while the processability of stainless steel powders had been investigated [8-9]. The role of laser sintering parameters, powder particle morphology, and powder oxygen content in the sintering response of laser sintered M3/2 and M2 high speed steel powders had been described [1-4].

The inference from these studies is that the transition from solid phase sintering to liquid phase sintering via full melting/solidification with increasing laser power at low scanning rates is the underlying mechanism guiding the SLS process. Marangoni flow in the melt pool was discovered to have affected the morphology of the sintered tracks at high laser power and low scanning rates. The agglomerate sizes increased with increasing laser power or decreasing scan rates. Moreover, it was observed that powder particle size less than 38µm had low sintered density as a result of high oxygen level.Corresponding author, Tel Direct (0113) 343 2359; Fax: (0113) 343 2384 E-mail address: r.f.cochrane@leeds.ac.uk (Dr R.F.Cochrane)It had been suggested that the surface tension gradient is a significant factor causing the spherodisation of the tracks or the break up of the liquid cylinder [1-4].

The energy exchange in the SLS of titanium powders via the application of the principle of the conventional coupling between laser and material as well as the phenomenon of the gray body behaviour of granulated media had been well elucidated [10]. Three key repetition domains, namely high repetition rate domain, lower domain and intermediate domain, were identified with respect to pulse energy and heating effects [10]. It was concluded that the intermediate domain constitutes the optimal processing window characterised by fast and localised heating which minimises the plasma effect [10]. Furthermore, the influence of laser parameters on the materials properties of laser sintered titanium and platinum-alloyed powders were investigated [11]. It was established that with the appropriate laser energy deposition in the powder bed, the degree of porosity, the density, as well as the crystalline microstructure can be controlled to meet the requirements of the finished work piece [11].

The influence of process parameters and heat treatment on the microstructure of direct laser fabricated TiAl alloy samples was investigated [12]. The findings from this study revealed that a variety of microstructures, ranging from cellular to columnar to dendritic morphologies, is obtainable upon the variation of the laser power and scanning speed. Furthermore, heterogeneous and finer microstructure were imparted to direct laser fabricated samples on comparison with the conventionally processed engineering parts made of titanium alloys. Subsequent heat treatment of the direct laser fabricated samples eventually resulted in the elimination of the compositional heterogeneity and coarseness of the samples [12].No systematic work on the effect of processing parameters on the microstructural development of aluminium alloys employing SLS has been published to date.

The aim of this work is to investigate the impact of the laser processing parameters on the densification and attendant microstructural characteristics of aluminium powder parts produced directly from the SLS process using a continuous wave CO2 laser beam. The microstructure of direct laser fabricated Al-12Si parts is characterised as a function of the process parameters.2. Experimental procedure.The starting material was inert gas atomized 45-75µm Al-12Si powder (spherical shaped particle), supplied by ALPOCO Limited. The loose apparent density, tap density, and flowability of the powder were determined according to the MPIF standard tests 01, 46 and 03 methods. A Philips XL30 ESEM Scanning Electron Microscope (SEM) with link Systems EDX and image capture accessories was employed to observe the particle shape. The chemical composition of the powder as supplied by ALPOCO is presented in Table 1.

The chemical composition of the powder as supplied by ALPOCO is presented in Table 1. Multiple layer samples were produced in Al-12Si powder by a Synrad 240W CO2 laser of 0.6 mm beam diameter having Gaussian profile. The laser processing parameters were used over the following ranges: laser power (100-200 W), scan rate (80-200 mms-1), scan spacing 0.1-0.3 mm, and layer thickness 0.25-1.00 mm. The description of the experimental apparatus and the procedure adopted for building multiple layer samples are given in [6]. The density of the fabricated samples was measured by using the volumetric method according to [6]. The processing condition for each sample was repeated at leasttwice and the density mean value recorded twice. Samples for metallographic examination were prepared and etched in dilute Keller’s reagent.Table 1: Elemental Composition of the Powdered Samples as Supplied by ALPOCO (Weight %).

Samples for metallographic examination were prepared and etched in dilute Keller’s reagent.Table 1: Elemental Composition of the Powdered Samples as Supplied by ALPOCO (Weight %). Powdered Samples Al-12SiAluminium SiliconOxygenIron Copper88.4012.10.360.083. Results.3.1: Powder Characterisation.Table 2: Powders’ Characterisation.Al-12SiLoose Apparent Density (gcm-3) 1.4347Tapped Density (gcm-3) 1.7238Flowability (s/50g)4.90Figure 1: The spherical particle shape of the Al-12Si powder.Table 2 and Figure 1 above indicate that both loose apparent and tapping densities of the powder are function of the powder’s characteristics such as particle shape, particle porousity and particle size distribution. A comparison of loose apparent and tapping densities shows that 20.15% percentage increase in density was obtained upon the tapping of Al-12Si powder.

A comparison of loose apparent and tapping densities shows that 20.15% percentage increase in density was obtained upon the tapping of Al-12Si powder. This is explained by the fact that its spherical particles are suspected to have packed without bridging or arching, thus creating empty spaces, as the particles move easily past each other because of their smooth surfaces.3.2: Effect of SLS Processing Parameters on the Sintered Density of Al-12Si Powder.Figures 2, 3, and 4 show the effects of the direct laser sintering process parameters on the sintered density of the Al-12Si powder. Generally, as shown in Figure 2, an increase in the laser scanning speed results in the reduction of the sintered density of Al-12Si powder over the range of laser power studied 100 W-200 W. A consideration of the relationship between the sintered density and the laser power of Al-12Si powder shows that sintered density increases as the laser power increases.

A consideration of the relationship between the sintered density and the laser power of Al-12Si powder shows that sintered density increases as the laser power increases. For instance, densities of 65.9%, 70.8%, and 80.2% were obtained when laser powers of 100 W, 150 W, and 200 W respectively were employed for SLS at the scanning rate of 120 mms-1 (Figure 2). Figure 2 also shows that the densities of Al-12Si sample sintered at 200 W (71.5% and 71.3%) are found to lie between the density values obtained at 100 W (70.4% and 66.3%) and 150 W (72.9% and 73.6%) for scanning rates of 80 and 100 mms-1. The normal expectation is that the density values obtained for 200 W over these scanning rates should be higher than those values obtained at both 100 W and 150 W because of the likelihood of the formation of greater quantity of liquid phase that promotes densification at higher laser power for which greater energy density is being deposited on the powder bed.

As the scanning rates increase over the range of 120 mms-1 to 200 mms-1, density values obtained for 200 W are indeed found to be higher than those obtained for 150 W and 100 W for corresponding scanning rates.

Furthermore, Figures 3 and 4 reveal that there exist inverse relationships between the sintered density and both scan spacing and layer thickness.100W (AL-5)200W (AL-5)150W (AL-5)Graph of % sintered density versus scanning rates.55.060.065.070.075.080.085.00.050.0100.0150.0200.0Scanning rates (mm/s) % Sintered densityFigure 2: Variation of sintered density of Al-12Si powder with the scan rates when laser power varies between 100 W -200 W and scan spacing and layer thickness were held fixed at 0.1mm and 0.25mm respectively.505560657075808500.10.20.30.40.5Scan spacing (mm)% Sintered density(a)5560657075808500.20.40.60.81Layer thickness (mm)% Sintered density(b)Figure 3: Variation of the sintered density of Al-12Si powder with (a) scan spacing and (b) layer thickness.3.3: Effect of SLS Processing Parameters on the Microstructure of Al-12Si Powder.Figure 4 shows the nature of the pore structures obtained in direct laser sintered Al-12Si specimens as a function of the processing conditions.

Although in each case, the microstructures contain small pores, the connectivity and orientation of pores are largely dependent on the choice of processing parameters. For instance, elongated pores which are nearly perpendicular to the building direction were found in a sample fabricated with a laser power of 200 W; scan rates of 120 mms-1; scan spacing of 0.1 mm and layer thickness of 0.25 mm (Figure 4 (a)). The eutectics are fully dense and consist of dendrites which are oriented nearly perpendicular to the build direction. Upon increasing the scan spacing to 0.5 mm, it was discovered that the orientation of the pores becomes undefined relative to the building direction (Figure 4 (b)). The eutectics obtained for the scan spacing of 0.5mm are observed not to have been as fully dense as the one obtained for the sample made with scan spacing of 0.1 mm.

The density was markedly higher (80.2%) for the sample made with the scan spacing of 0.1 mm than that obtained for the sample made with a scan spacing of 0.5 mm (53.7%). The findings so far indicate that increasing the scan spacing leads to an increase in pore sizes, less dense eutectics and less oriented pores. Furthermore, similar results were obtained with increasing layer thickness (Figure 4c), increasing scan rates (Figure 4d), and decreasing laser power (Figure 4e).Figures 5, 6, and 7 show optical micrographs of laser sintered Al-12Si powder processed at energy densities of 100 Jmm-3, 67 Jmm-3, and 50 Jmm-3 respectively. Black regions in the micrographs are an indication of the presence of pores while the grey region corresponds to the sintered mass.

Black regions in the micrographs are an indication of the presence of pores while the grey region corresponds to the sintered mass. The pores are found to be responsible for a significant decrease in the densities of the sintered parts.Figure 5: Sections through the micrograph of AL-5 (200 W-80 mms-1) generated by 100 Jmm-3.Under these conditions, a coarse eutectic microstructure is observed as shown in Figures 8(a) and (b) for samples fabricated with 100Jmm-3. The microstructure contains small pores (Figure 5). Furthermore no continuous network of pores was observed; rather, irregular shaped pores with no orientation to the building direction, surrounded by a fully dense aluminium-silicon matrix are formed. As the energy density is reduced to 67 Jmm-3 (Figure 6), consequently, a finer microstructure as shown in Figures 8(c) and (d) results from the processing condition.

As the energy density is reduced to 67 Jmm-3 (Figure 6), consequently, a finer microstructure as shown in Figures 8(c) and (d) results from the processing condition. Here, the finer dendritic microstructure and the interdendritic porosity found could be attributed to the shrinkage which occurs during solidification and/or with the presence of some entrapped argon. As the energy density is reduced to 50 Jmm-3, (Figure 7) it can be seen that the heat input is just enough to sinter the powder. The consequence of this is that the dendritic microstructure can be seen to have become still finer but the interconnectivity of the porosity increases. Figure 8 reveals that, for all samples, the dendrites become coarser and porosity increases towards the bottom part of the sintered samples.4.0: Discussion:The most important parameter that affects the density and microstructure of direct laser sintered materials is the energy density.

Energy density during the laser sintering process is defined by the following relationship:E = P / (V) x (d) x (h)……………………….(1)Where E is the incident laser energy density (Jmm-3), P is the laser power (W), V is the scanning rate (mms-1) d is the scan spacing (mm) and h is the layer thickness (mm). The relationship above shows that the energy density for the laser sintering process increaseswith increasing laser power, and decreasing scanning rates, scan spacing and layer thickness.1Figure 6: Sections through the micrograph of AL-5 (200 W-120 mms-1) generated by 67 Jmm-3Figure 7: Sections through the micrograph of AL-5 (150 W-120 mms-1) generated by 50 Jmm-3.Therefore, plausible explanation for the observation noted in the values of sintered density when employing varying scanning rates and laser power (Figure 2) could be attributed to the optimisation of the energy density applied to the sintering process.

As the energy density dissipated to the Al-12Si powder bed increases, the amount of liquid phase available for the densification mechanism increases until a critical value of energy density is reached. After the attainment of the critical energy density, further increament in the applied energy density results in spherodisation due to the oxidation of the Al-12Si powder particles. Moreover, spherodisation occurs owing to longer powders’ particles exposure at high powder bed temperature. Spherodisation inhibits interparticulate bonding across the layers, thereby promoting the formation of porosity. Therefore, since spherodisation has crept in, the sintered densities of the samples are less than thatfor a longer duration. This occurs after the energy density has been optimised.

Therefore, since spherodisation has crept in, the sintered densities of the samples are less than thatfor a longer duration. This occurs after the energy density has been optimised. Moreover, insufficient densification observed in the microstructures of direct SLS of Al-12Si powder below the critical value can be explained on the basis of the laser energy density reducing the amount of liquid phase available for sintering. This results in increased porosity and a reduction in the sample density.50.055.060.065.070.075.080.085.00.020.040.060.080.0100.0Energy Input per Unit Volume (J/mm-3)% Sintered Density.Figure 9: Graph of % apparent density versus the energy input per unit volume for Al- 12Si aluminium powders for all processing conditions investigated.The direct SLS technique employs a very high power density laser beam which is capable of generating typical cooling rates of the order of 103 to 104 K/s [13-14].

The fine microstructure observed in direct laser sintered Al-12Si sample processed with 67 J/mm-3 is a consequence of this high cooling rate. The implication of this is that the nature of the microstructure observed when varying laser power, scanning rates, scan spacing and layer thickness is dependent on the duration of the interaction between Al-12Si powder and the laser beam. The relationship between energy density and duration of laser- material interaction defines the operational regime of the direct SLS process which results in a unique temperature gradient (G), solidification rate (R) and cooling rate (T = G.R). It was noted that increasing the ratio G/R results in gradual change in the solidification regime from dendritic to cellular dendritic and eventually to planar front growth [13]. Furthermore, a higher cooling rate (G.R) leads to greater undercooling thus producing more nuclei of finer grains.

Furthermore, a higher cooling rate (G.R) leads to greater undercooling thus producing more nuclei of finer grains. Hence, the parameters (G/R) and (G.R) control the microstructural features and determine the scale of the microstructure [14-16]. With regard to laser sintered Al-12Si samples processed at 100 Jmm-3, a longer interaction time encountered by Al-12Si powder due to the combination of lower scanning rate (80 mms- 1) and higher laser power (200 W) leads to the delay in the initiation of solidification process. Consequently, the temperature of the substrate material rises and the temperature gradient of the interface and the cooling rate are reduced. This leads to the formation ofcoarse dendritic microstructure (Figures 8 (a) and (b)).

This leads to the formation ofcoarse dendritic microstructure (Figures 8 (a) and (b)). Considering the case of Al-12Si sample processed at the critical energy density of 67 Jmm-3, made up of higher scanning rate (120 mms-1) and high laser power (200 W), it is speculated that superheating is not as high and the average temperature gradient at the interface is high thereby leading to faster cooling rate which results in a finer microstructure (Figure 8 (c) and (d)). As the deposited energy density (50 Jmm-3) is less than the critical value, the heat input is just sufficient to sinter Al-12Si powder particles which eventually results in higher surface tensions of the melt pool and formation of cylindrical geometry.

The contact area between two such layers will be smaller thereby leading to poor heat conduction, low cooling rate and dendritic microstructure (Figure 8 (e) and (f)).During laser sintering, cooling of the molten pool occurs via the substrate and the surrounding atmosphere. Initially, heat loss through the Al-12Si substrate layer 1 results in rapid cooling via the substrate rather than through convection and radiation. The consequence of this is the directional growth of the dendrites counter to the cooling direction. A careful observation of Figure 10 shows that it seems that the tops of the grains in the previous layer 1 are partially re-melted and then serve as nuclei leading to epitaxial growth counter to the cooling direction. This process is repeated for many consecutive layers until the heat loss via the substrate is no longer dominant during the solidification mechanism.

This process is repeated for many consecutive layers until the heat loss via the substrate is no longer dominant during the solidification mechanism. Consequently, the addition of a new layer 2 of material to the pre-existing layer results in the solidification of the region on top of the pre-existing layer 1 upon re-heating, thus grain growth takes place in pre-existing solid layers of the sintered sample (Figure 10). As the sample height increases, heat conduction to the substrate reduces and the build temperature increases. Therefore, the temperature change as a consequence of re-heating in the region underneath the interface affects a larger part of the build. As a result, the microstructure in the bottom part of the build (Figures 8) becomes coarse and the differences between re-heating and solidification become insignificant because heat extraction to the substrate for the lower region is unaffected.

The disappearance of layer bands in the Figures 5-7 when the energy density varies between 50-100 Jmm-3 could be attributed to the same reason.4.0 Conclusions(1) Laser power, scanning rates, scan spacing and layer thickness are found to control the densification and the resultant microstructural characteristics of the laser sintered Al-12Si parts.(2) Sintered density increased as the energy density until a critical energy input per unit volume (67 Jmm-3) was realised for which the optimum sintered density (80.2%) was obtained.

Above the critical energy density a gradual reduction in density was noted to have occurred until a point is reached at which the sintered density remains unchanged with further increases in energy density.(3) Parts made with an energy density of 50 Jmm-3 were characterised by interconnected porosity and a fine dendritic microstructure while parts fabricated with an energy densityof 100 Jmm-3 consists of a coarse dendritic microstructure with discontinuous, irregularly shaped pores surrounded by a fully dense aluminium-silicon matrix.BABLayer 3ABLayer 2BALayer 1Figure 10: An enlargement of point 1 on the Figure 6 showing various points on the melt- back could be clearly observable across the layers (A) and areas where there had been no melt back resulting in porosity (B) during laser sintering.References[1] Niu, H. J. and I. T. H. Chang (1998).

Microstructural and chemical analyses were conducted using opticalmicroscopy, scanning electron microscopy and energy differential system. It is worthnoting that the element Al loss rate was about 8% under different process parameters.As a result, the microstructures of EBSM Ti47Al2Cr2Nb samples were composed ofcolumnar β grains, α/α2and α2/γ lamellar. Tensile tests were carried out to understandthe mechanical properties to the corresponding microstructures. Ultimate TensileStress (UTS)at room temperature is much lower than that at a high temperature.IntroductionIntermetallic TiAl alloys are gaining more and more attentions because of theoutstanding properties such as the low density(3.9-4.2g.cm-3), high strength andhigh-temperature oxidation and creep resistance.

TiAl alloys appear as promisinghigh-temperature materials used in automotive, energetic and aerospace industries.However, TiAl alloys exhibit poor processing plasticity at room temperature which500leads to poor workability at room temperature. Recently, electron beam selectivemelting (EBSM) has been studied as a feasible method for manufacturing Ti6Al4Vand Ti-Al intermetallic [1-3].Electron beam selective melting is one of the additivemanufacturing methods with highfabrication quality. Comparedtomanufacturing [4], electron beam selective melting has higher energy density and scanspeed which reduce the time and cost of manufacturing. In addition, EBSMtechnology has advantages of vacuum fabrication environment in fabricating impuritysensitive and highly reflective materials.

In addition, EBSMtechnology has advantages of vacuum fabrication environment in fabricating impuritysensitive and highly reflective materials. Therefore it is of particular interest toinvestigate the EBSM fabrication of intermetallic TiAl components.TiAl alloys have four typical microstructures and the mechanical properties ofTiAl alloys are strongly dependent on the microstructure. In electron beam selectivemelting, columnar crystals trend to be formed because of the high cooling rate and bigtemperature gradient.

In electron beam selectivemelting, columnar crystals trend to be formed because of the high cooling rate and bigtemperature gradient. The scan pattern has an effect on the temperature gradient,which would further influence the growth of grains and the mechanical properties.This work investigated the microstructures and resultant mechanical properties ofsamples fabricated employing three different scan patterns, namely S-shaped scan line,Z-shaped scan line and interlayer orthogonal S-shaped scan line (also called interlayercross scan line).ExperimentsIn this study, square samples were fabricated by electron beam selective meltingfrom pre-alloyed Ti47Al2Cr2Nb powders with particle sizes between 45 and 150μm(Fig.1). The powder had a chemical composition (in atomic percentage)of46.51%aluminum, 0.02% niobium, 0.02% chromium, balance titanium.

The powder had a chemical composition (in atomic percentage)of46.51%aluminum, 0.02% niobium, 0.02% chromium, balance titanium. The powder exhibitsfine surface and highly spherical shape, and little satellite particles are also visible.501laserFig.1microscopic view of Ti47Al2Cr2Nb powderFig.2 shows the schematic diagram of EBSM-250 system designed by TsinghuaUniversity. Compared with commercial EBSM equipment, EBSM-250 system canmix two kinds of powder together and then fabricate functionally gradient materials.In the system, electron beam is generated in the electron gun and accelerated at avoltage of 60KV.

The electron beam is focused by focusing coil and scanned bydeflection coil to preheat and selectively melt powders layer-by-layer.FilamentCathode70kVGate (0~-1500V to the cathode)Electron beam gunAnodeFocusing coilDeflection coilVacuum chamberPowder storageElectron beamPowder spreaderPartsVertical pistonFig.2 schematic diagram of EBSM-250 system designed by Tsinghua UniversityIn this work, a preheating beam current of 20mA, a melting beam current of 6mAa hatching distance of 0.1mm and a scan speed of 0.2m/s were employed.Components were built directly on 316L stainless steel base plate dimensioning10mm in thickness and 90mm×90mm in area. Before fabrication, the stainless steelbase plate was preheated to about 800(cid:263).The effect of scan pattern on microstructureand mechanical properties were investigated systematically by varying the scan502pattern.

Flake samples were taken for tensile tests at different temperatures on auniversal material tension machine.Results and discussion3.1 Aluminum lossAs reported before, element Al loss is dependent on energy input and beam current.High energy input and beam current will result in high loss rate. CarolinKorner[5]found that line energy of 1.1J/mm and beam current of 9.2mA cause a Al lossof about 4% while line energy of 0.7J/mm and beam current of 3.5mA reduce the lossto about 1at.%.Denis Cormier et al[6] fabricated Ti-47Al-2Cr-2Nb samples employinga scan speed of 0.1m/s and a beam current of 13mA, and found approximately 6.1%503of the aluminum vaporized during the process. In this paper, specimens werefabricated employing a beam current of 6mA, a scan speed of 0.1m/s and a layerthickness of 0.1mm underthree different scan patterns.

%L→β+L→β+α→α→α2→lamellar (α2+γ)(2)42at. %< Al<50at. %L→β+L→β+α+L→β+α+γ→α+γ(3)42at. %< Al<50at. %L→α+L→γPreviousinvestigationsrevealthattheprocessingparametershaveremarkableinfluence on Al loss. Further investigation is needed to adjust the energyinput to avoid over-heating and reduce Al evaporation.3.2 Microstructure evolutionMicrostructure evolution is closely related to the cooling and heating historyduring the process. EBSM process has a characterization of rapid solidification and504arepeated heat cycling. In general,the microstructures of EBSM fabricatedTi47Al2Cr2Nb alloy consist of highly columnar prior β grain along the buildingdirection and α2 plates.

Many researchers [7-9] reported such columnar microstructures in theadditive laser manufacturing (ALM) and EBSM of Ti6Al4V and Nickel-base superalloy. Components fabricated by EBSM are surrounded by loose powders which actas thermal insulation while the base plate and deposited materials act as heatdissipation parts. Due to this gradient, a preferred orientation of the dendrites in theforms of arrays with a typical spacing is obtained. For β grains, a body-centered cubiccrystal, the [100] is the preferential crystallization direction. The elongated grainsgrow towards the molten pool and tilt a certain angle to the Z-direction. As shown inFig.5,the EBSD pattern results also reveal the columnar preferred orientation.

The elongated grainsgrow towards the molten pool and tilt a certain angle to the Z-direction. As shown inFig.5,the EBSD pattern results also reveal the columnar preferred orientation. As theinterlayer cross S-shaped scan line leads to the deviation of the direction of thetemperature gradient, the growth of columnar crystal were interrupted.Fig.4(a) showsthe optical microscopy of Cross-shaped scan specimen, the main structure is stillcolumnar grains which is parallel to the building direction. Compared to the S-shapedscan specimen, the size of columnarβ grainof Cross-shaped decrease from 400μm to~200μm and the grains become more equiaxed.507Fig.5 Orientation image and Pole figure of Cross-scan Ti47Al2Cr2Nb alloyFig.6 (a), (c), (e) show the microstructures obtained using the three different scanpattern. A thin layer ~1mm of martensite α phase is observed at the top of the sampleas a result of high cooling rates. Qian et al [10]and S.S.

A thin layer ~1mm of martensite α phase is observed at the top of the sampleas a result of high cooling rates. Qian et al [10]and S.S. AL-BERMANI [11]reported thatthe cooling rate of top layer from liquid was ~7×104K·s-1 and 103~105K·s-1 in SLMand EBM Ti6Al4V. Although the exact cooling rate of Ti47Al2Cr2Nb is unknown,cooling rate of the new deposited layer is high enough to get martensite α. As samplesincrease in height, the temperature of the deposited layers are kept in relative hightemperature and transformation α→α2, α2→α2+γ take place. Fig6 (b), (d), (f) show thebasket weave/ Widmanstatten α/α2 laths and α2/γ lamellar in the middle of the samples.In the middle and bottom of the samples coarsen α can be observed with width of~5μm. The form of coarse α lath is influenced by thermal cycling during the newdeposited layers.

The form of coarse α lath is influenced by thermal cycling during the newdeposited layers. Deposited part of the samples was kept in a relatively hightemperature, for SLM Ti6Al4V the temperature is508(a) top of Cross-scan (b)middle of cross scan(c)top of S-scan(d)middle of S-scan(e) top of Z-scan(f) middle of Z-scanFig.6 SEM photos of samples3.3 Tensile TestTensile specimens were tested on an Instron-5967 tensile machine using specialgrips at an engineering strain rate of 2×10-4 s-1.Tensile properties at room temperatureand the high temperature of 700(cid:263)are illustrated in Fig.7. Fig.8 shows the fracturesurface features for the S-shaped scan specimens at room temperature and the hightemperature. Translamellar fracture appears to be the major failure mode in the two509samples at different temperatures. The UTS at 700(cid:263)is dramatically enhanced.

Translamellar fracture appears to be the major failure mode in the two509samples at different temperatures. The UTS at 700(cid:263)is dramatically enhanced. Thetensile strength at the room temperature of Cross-scan and S-scan samples is271.3MPa and 362.6MPa, while the tensile strength at the high temperature is679MPaand769.6MPa respectively. The reason may be that specimens for high temperaturetest experienced a short time(5min) of heat preservation process which has a certaineffect on reducing the residual stress. During the additive manufacturing process,repeated heating and cooling cycles in the material induce important thermal gradientsand then may cause residual stress in samples after process. Many researchers [12-13]investigated the residual stress in selective laser melting process and the results showthat the main residual stress is tensile stress.

Many researchers [12-13]investigated the residual stress in selective laser melting process and the results showthat the main residual stress is tensile stress. The exist of tensile stress in thespecimens may decrease room temperature tensile strength.800Cross-scanS-scan700 oC) a P M600700 oC(s s e r t S g n i r e e n i g n E400200RTRT00123Engineering Strain (%)Fig.7 the tensile curves of different samples at different temperatures510(a) Room Temperature(b) 700(cid:263)(cid:3)Fig.8S-shaped scan sample fracture surface featuresConclusionsTi47Al2Cr2Nb alloy square samples were fabricated by electron beam selectivemelting from pre-alloyed powders. The scan pattern has an effect on the temperaturegradient, which would further influence the growth of grains and the mechanicalproperties.

The scan pattern has an effect on the temperaturegradient, which would further influence the growth of grains and the mechanicalproperties. In order to study the effect of electron beam scan pattern on themicrostructure evolution, three different scan patterns were employed: S-shaped scanline, Z-shaped scan line and interlayer orthogonal S-shaped scan line.The element analyses by EDS demonstrated that the element loss rate of Alranged from 6at. % to 8at. %employing different scan patterns. According to the Ti-Altwo-phase diagram, the microstructures of EBSM Ti47Al2Cr2Nb samples werecomposed of columnar β grains, α/α2 and α2/γ lamellar. The dendritic crystal growthdevelops along the direction of heat dissipation, which is influenced by scan patterns.As the interlayer cross S-shaped scan line leads to the deviation of the direction of thetemperature gradient, the growth of columnar crystal were interrupted.

Therefore thesize of the columnar crystal in cross-scanned specimens appeared smaller than theothers.Tensile tests at room temperature and high temperature of 700(cid:263) were carriedout to understand the mechanical properties to the corresponding microstructures. Atthe room temperature, the specimens employing interlayer cross S-shaped scan lineachieve an ultimate tensile stress (UTS) of 271.3MPa, while the UTS of the511specimens employing S-shaped scan line is 362.6MPa.It is worth noting thatUTS atthe temperature of 700(cid:263) is much higher, 679MPa and769.6MPa for cross scan andS-shaped scan respectively.AcknowledgementsThis work has been supported by the funding of 2013 Beijing Science andTechnology Development Project, project numbers D13110400300000 andD131100003013002.References[1] Denis Cormier, Ola Harrysson, Tushar Mahale, Harvey West.

Metallurgical and Materials Transaction A, Volume 41A, December 2010.[12] B. Vrancken, R. Wauthle, J.-P.Kruth, J.Van Humbeeck. Study of theinfluence of material properties on residual stress in selective laser melting.[13] Peter Merceies and Fean-Pierre Kruth. Residual stresses in selective lasersintering and selective laser melting. Rapid Prototyping Journal, 12/5(2006) 254-265.513.

fect of Scan Pattern on the Microstructural Evolution of Inconel 625 during Selective Laser MeltingMd Ashabul Anam, JJS Dilip, Deepankar Pal, Brent Stucker Department of Industrial Engineering J.B Speed School of Engineering, University of Louisville, Louisville, KY 20292Abstract:Selective laser melting (SLM) involves highly localized heat input and directional solidification, which enables novel microstructure control through the development of scanning strategies and related process variables. A careful study of scan pattern is important to understand microstructural evolution during SLM. In this study, various types of scanning strategies were used to build samples of Inconel 625. Microstructure differences due to different scan patterns in as-built Inconel 625 samples were then studied in detail. Microstructure samples showed grains with cellular substructure with enriched regions of Nb and Mo in the inter arm spacing.

Microstructure samples showed grains with cellular substructure with enriched regions of Nb and Mo in the inter arm spacing. The grains were observed to grow preferentially in the build direction, but there were also clear effects of grain orientation differences due to scan direction effects.1. Introduction:Selective Laser Melting (SLM) is characterized by highly localized high heat input for very short periods of time, which has a profound impact on the microstructure. SLM provides a wide range of advantages over conventional manufacturing techniques, but high temperature gradients and rapid solidification cause high thermal stress build-up and the presence of non-equilibrium phases.

Orientation of grains is largely controlled by the heat conduction direction, thus laser scanning strategy becomes a powerful tool for control of grain orientation, and hence the microstructural texture.This paper focuses on the analysis of different scan strategies in order to obtain insight into the development of microstructural texture in Inconel 625 during SLM processing. It has been observed that SLM is a complicated process where the evolution of the microstructure depends on a large number of parameters such as laser power, scan speed, layer thickness, and scan strategy [1]. Most researchers have concentrated on the influence of process parameters on mechanical properties such as relative density, hardness and fatigue.

Most researchers have concentrated on the influence of process parameters on mechanical properties such as relative density, hardness and fatigue. From those studies, 20 µm layer thickness was found to be optimal for In625 and for laser power and scan speed, a range of parameters were found the purpose of microstructural analysis, 20 µm layer thickness and a 195W-800mm/s laser power-scan speed combination were primarily used. In this research two types of scanning strategies were investigated to understand their effects on the microstructures of IN625.to provide good mechanical properties. For2. Scan Patterns in SLMSLM machines typically utilize one of 2 basic scan strategies for the interior sections of parts. These strategies are either to scan the bulk of the part utilizing parallel stripes or by scanning the bulk of the part utilizing a block pattern.2.1. Rotated Stripe PatternA schematic of a rotated stripe pattern is shown in figure 1.

Rotated Stripe PatternA schematic of a rotated stripe pattern is shown in figure 1. When using a stripe pattern, each layer is divided into a series of parallel stripes (red dashed line in figure 1) that run across the entire length of the cross-section of any closed contour within a layer. Within each stripe, raster scan vectors are used. During every new layer, the stripes rotate counterclockwise by ~67ᵒ compared to the previous layer, creating a crosshatch pattern. Samples were cut from cylindrical fatigue specimens of diameter 10 mm and height 85 mm which were built in the Z direction in an EOS M270 machine using In625 powder from EOS (figure 2a). Both horizontal and vertical sections were taken and analysis was done using optical microscopy as well as scanning electron microscopy.Layer #3Build Direction (Z)~67ᵒLayer #2Layer #1Figure 1: Schematic of a stripe scan pattern showing raster scan vectors (arrows) within each stripe (red dashed line).

For every new layer, the stripes rotate by 67ᵒ compared to the previous layer.a.364a.b.(Y)67ᵒ(X)Figure 2: (a) Samples were cut from vertically built (Z orientation) cylindrical specimens. (b) Microstructural evidence of the stripe pattern, from a horizontal (XY plane) cross-section. Since the view plane is not perfectly horizontal, multiple layers intersecting the plane can be seen. The stripes rotate 67ᵒ for each new layer thus create a crosshatch patterns.(Z)(Y)Figure 3: Melt-pool arrangement as seen from a vertical cross-section (YZ plane) for a stripe pattern.In625 samples built using a rotated stripe scan strategy were etched and observed using optical microscopy. Figure 2(b) shows a vertical cross section of the sample where laser tracks are clearly distinguishable. The view plane in figure 2(b) is not perfectly horizontal (tilted slightly) which enables one to view multiple layers in the same image.

The view plane in figure 2(b) is not perfectly horizontal (tilted slightly) which enables one to view multiple layers in the same image. The angle between laser tracks of consecutive layers was found to be 67ᵒ as illustrated in figure 1. Two stripes in the same layer, as well as their overlap, are visible in the top left corner of figure 2(b). In figure 3, the melt-pool arrangement in the vertical cross-section (YZ plane) is shown. It was observed that the shapes ofthe melt-pools are quite similar (Gaussian) but their sizes are different. This is because scan vectors from different layers are not parallel to each other and not perpendicular to the view plane as well. Thus, different oblique sections of the melt-pools show different sizes.2.2. Alternating Block PatternSchematics illustrating an alternating block scan strategy are shown in figure 4. In each layer, rectangular blocks with raster scan vectors are used and the scan vector rotates 90ᵒ between adjacent blocks.

In each layer, rectangular blocks with raster scan vectors are used and the scan vector rotates 90ᵒ between adjacent blocks. Between layers, the block shifts half of its width creating a “brick-like” stacking sequence. This scenario is clearly depicted in figure 4b where a schematic of a vertical cross- section (XZ plane) is shown (black circles represent scan vectors in the Y or –Y direction and black rectangles represent vectors in the X or –X direction). A block pattern such as this is typically designed to replicate the 0-90-0-90 stacking sequence often used in composites. However, the region labeled “Section A” in Figure 4b shows a stacking-based anomaly where a portion of that region has a 0-90-0-90 repeating structure and another portion of that same region has a 0-0-0-0 stacking sequence.

Thin rectangular samples 2 mm thick, 40 mm long and 5 mm high were fabricated using In625 (figure 5) in an EOS M270 machine and analysis was done using optical microscopy as well as scanning electron microscopy to investigate this phenomena.a.b.Layer #3Build Direction (Z)Layer #2Section ABlock WidthLayer #1XFigure 4: (a) Schematic of a block scan pattern showing raster scan vectors (arrows) within each block, (b) Vertical cross-section (XZ plane) showing scan line patterns between layers. In each new layer the block shifts half of its width.Layer #3Layer #2Layer #1Section AFigure 5: 40 mm x 5 mm x 2 mm blocks. Section A from the sample (Right) was cut and analyzed to understand the microstructural effects of the block scan strategy.Section A from the sample in figure 5 was cut and polished and both vertical (XZ plane) and horizontal (XY plane) cross-sections were observed for microstructural features.

Figure 6 shows the etched vertical cross section where melt-pool arrangement is clearly visible. The image can be separated into two distinct areas which were formed due to different scan vector overlaps. The left section of the image has a unidirectional melt-pool which were formed by blocks where all the scan vectors are parallel to each other for every layer (the direction was perpendicular to the view plane). The right section of the image shows scan vectors where layers create alternating hatch patterns. Section A of figures 4b and 5 illustrate how the pattern shown in figure 6 was created.2 different layers with scan vectors parallel to each other.2 different layers with scan vectors perpendicular to each other.Build Dir. (Z)(X)Figure 6: Melt-pool arrangement in a vertical cross-section (XZ plane) using a block pattern scan strategy.(Y)(X)Figure 7: Block pattern, horizontal cross-section (XY plane) showing laser tracks for a block scan pattern.

Since the view plane is not perfectly horizontal, multiple layers of laser tracks are seen.3. Effect of Scan Pattern on MicrostructureSLM is characterized by highly localized heat input for a very short period of time leading to rapid solidification as the melt pool undergoes liquid to solid transformations. This may result in formation of non-equilibrium phases and changes in general microstructural features. Finer cellular dendritic structures evolve due to high cooling rates compared to conventional manufacturing processes [2, 3, 8-11]. In addition, cellular/dendritic structures, and thus grain structures, are also controlled by microstructures formed at the boundaries of previously solidified layers (figure 8).

In addition, cellular/dendritic structures, and thus grain structures, are also controlled by microstructures formed at the boundaries of previously solidified layers (figure 8). Although the mechanism of grain growth is highly complex and influenced by many variables, local heat conduction direction plays an important role in determining the orientation of grains.Temperature gradient is the highest along the midline of the scan track and at the back of the moving melt pool, and decreases radially [5]. Thus grains tend to grow in this direction (figure 8). However, due to overlap, major portions of laser tracks and their grains remelt during subsequent scanning. Due to the fast cooling conditions, the grains grow epitaxially toward the center of the new melt pool [4,6,7]. Based upon these conditions, the orientation of grains is thus highly dependent on the scanning velocity and scanning strategy.

Based upon these conditions, the orientation of grains is thus highly dependent on the scanning velocity and scanning strategy. Consequently, the scanning strategy can become a powerful tool to control the grain orientation, and hence the microstructural texture [7,16,19,20].(a)(b)Figure 8: (a) SEM image of a horizontal (XY) section showing cellular/dendritic growth (red arrows) towards the center of the scan track (white arrow), (b) vertical (YZ) section showing radial growth of cellular dendrites. Clear melt-pool boundaries are observed which has an effect on grain size and orientation.Thijs et. al. (2010) studied different scan strategies for Ti-6Al-4V and found that grains grow preferentially in the build direction. They reported that elongated grains are parallel to each other and are tilted at an angle 19ᵒ away from building direction for a unidirectional scan pattern.

They reported that elongated grains are parallel to each other and are tilted at an angle 19ᵒ away from building direction for a unidirectional scan pattern. For bidirectional and grid-like scan patterns (figures 6 and 7), elongated grains were found to grow preferentially in the build direction (figure 9b).(a)(b)Figure 9: Orientation of elongated grains scanned with: (a) unidirectional scan vectors, and (b) alternating hatch scan vectors. View plane: Side view (YZ plane), Material: Ti-6Al-4V. [7]Similar phenomena were observed for In625. For a stripe scan strategy which rotates 67ᵒ for every new layer, more refinement of cellular/dendritic structures occurs due to maximum overlap in the melt, creating more elongated grains.

For a stripe scan strategy which rotates 67ᵒ for every new layer, more refinement of cellular/dendritic structures occurs due to maximum overlap in the melt, creating more elongated grains. Figure 10 shows optical microscopy of elongated grains oriented in the build direction (Z).ZFigure 10: Orientation of elongated grains (dark and white regions) formed due to the stripe pattern, which grow preferentially in the build direction (z) indicated by black arrows across multiple layers.Despite epitaxial solidification, melt pools are still distinguishable due to the appearance of dark bands in optical microscopy (white bands in SEM) (figure 11). These dark bands are more noticeable when more energy is applied. The dark contrast which distinguishes these bands arises due to enhanced precipitation of Nb and Mo which are preferentially etched [2,3,12-15].

The dark contrast which distinguishes these bands arises due to enhanced precipitation of Nb and Mo which are preferentially etched [2,3,12-15]. EDS analysis was done for this identification and is shown later in this paper.a .b .ZYXc .d .ZYXFigure 11: (a) Optical microscopy showing a vertical section of In625 where melt pools are distinguishable due to the appearance of dark bands. (b) Horizontal section showing dark band region. (c) SEM image showing the same section as ‘a’ but magnified to illustrate that SEM white regions represent optical microscopy dark bands. (d) High magnification SEM image of a horizontal section showing the cross–section of a single dendrite where the white region is formed due to segregation of Nb and Mo.Figure 12(a) shows a typical polished and etched SEM micrograph of the vertical cross-section of an In625 part built using a rotated stripe pattern.

It can be observed from the micrograph that the grains have grown epitaxially and there is grain continuity across multiple melt pools in the build direction. The columnar grains show cellular and dendritic sub-structure. It can also be observed that there is noticeable segregation (bright regions) within the intercellular and inter- dendritic arm spacing. Intercellular spacing is less than 1 μm (figure 13) which provides excellent strength and hardness that can be achieved both in the as processed and aged conditions.Figure 12(b) shows a typical polished and etched SEM micrograph of the vertical cross-section of an In625 SLM sample using a block scan pattern. The microstructural features show epitaxial grain growth from the melt. The grains show columnar morphology with a cellular / dendritic sub-structure. The individual cell/dendrite boundaries are illuminated by segregation of elements during solidification into the inter cell/ dendritic arm spacing.

The individual cell/dendrite boundaries are illuminated by segregation of elements during solidification into the inter cell/ dendritic arm spacing. There is also grain continuity across multiple melt pools, which is also common for multipass welds of this alloy. For alternating scan patterns, the elongated columns were found to have shorter length compared to rotated stripe patterns.Melt-pool boundaryInter-meltpool coarse structureMelt-pool boundarya.b .Z ZZScan DirectionFigure 12: SEM image showing vertical cross-section (YZ plane) of In625 sample built using (a) rotated stripe pattern and (b) alternating block pattern. Arrows in the right section in ‘b’ show alternating scan directions.Another distinguishable feature is that regions of coarse columnar structure appear in between melt pools in figure 12(b). These regions are less refined. Refinement in columnar structure is done by laser re-melting which is reported by several authors [10,17].

These regions are less refined. Refinement in columnar structure is done by laser re-melting which is reported by several authors [10,17]. Melt-pool depth for this experiment is approximately 60 μm, which influences three layers. Therefore overlap and remelting become a common phenomenon in SLM processing. However alternating block patterns leaves a significant amount of inter-meltpool regions which are not overlapped. These inter-meltpool regions undergo incipient melting, developing a continuous network of undesirable interdendritic phase. For rotated stripe patterns, higher overlap and thus finer cellular structure is found. Cellular structure size and direction have significant effect on material strength.

For rotated stripe patterns, higher overlap and thus finer cellular structure is found. Cellular structure size and direction have significant effect on material strength. Finer structures provide higher hardness and good fatigue strength [18].a .b .ZYXFigure 13: SEM image showing intercellular spacing less than 1 μm for a rotated stripe scan pattern in (a) vertical and (b) horizontal sections, which provide excellent strength and hardness.4. EDS AnalysisLocal chemical composition of the cell/dendrites and within inter cell / dendritic arms was evaluated qualitatively by using energy spectroscopy in SEM. The inter cell / dendritic regions show regions enriched in Nb and Mo.BAPosition Point A: On Matrix Point B: On Column BoundaryMaterial Composition (Wt. %) Nb Mo Cr 2.04 5.62 18.06Ni+Others 74.293.34 6.31 17.8272.53Figure 14: EDS spectrum comparison for two different points showing enriched Nb and Mo in dendritic arms boundaries.5.

%) Nb Mo Cr 2.04 5.62 18.06Ni+Others 74.293.34 6.31 17.8272.53Figure 14: EDS spectrum comparison for two different points showing enriched Nb and Mo in dendritic arms boundaries.5. Backscattered Electron ImagingThe samples were also observed in the BSE mode in SEM. The BSE imaging mode gives the atomic number contrast of elements in the microstructure under observation. The BSE SEM micrograph of the vertical cross section further confirms the segregation of elements in the inter cell/dendritic arm regions.Figure 15: BSE image of a vertical (YZ) cross-section at two different magnifications.6. SummaryTwo types of scan patterns were analyzed in detail using schematics as well as optical and SEM microscopy for In625 samples fabricated using selective laser melting. A rotated stripe pattern which uses raster scan vectors and which rotates ~67ᵒ for every new layer creates a crosshatch pattern.

A rotated stripe pattern which uses raster scan vectors and which rotates ~67ᵒ for every new layer creates a crosshatch pattern. A block pattern, on the other hand, uses raster scan vectors in each block and shifts half of its width between each new layer, thus creating brick-like building blocks and regions where no shift in scan vectors occur.In SLM, local thermal gradients inside the melt pool are the most important factor for determining microstructure. The heat conduction direction, based upon scan pattern, plays an in a important role cellular/dendritic fashion with the growth direction preferably oriented towards the center of the melt pool. However, due to overlap, major parts of each solidified track remelts and the grains grow epitaxially towards the center of the new, adjacent melt pool. Microstructures of previously solidified layers also control the direction of cellular/dendritic growth.in determiningthe orientation of grains.

Microstructures of previously solidified layers also control the direction of cellular/dendritic growth.in determiningthe orientation of grains. Solidification occursNoticeable amounts of segregation within the intercellular and interdendritic arms spacing were observed. Intercellular spacing is less than 1 μm, which contributes to the excellent strength and hardness that can be achieved both in as-processed and aged conditions. These cells are not individual grains but constitute a substructure inside one grain.Grains grow epitaxially and there is grain continuity across multiple melt pools in the build direction. However, for an alternating block pattern, the elongated columns are found to have shorter length compared to a rotated stripe pattern.Difference in cellular structure were observed in alternating block patterns.

Inter-meltpool regions undergo incipient melting, developing a continuous network of undesirable interdendritic phase, thus resulting in coarse cellular structures which should affect hardness and strength significantly.AcknowledgementThe authors would like to thank the staff at the University of Louisville Rapid Prototyping Center for help with fabrication of test coupons. The authors also acknowledge support of the National Institute of Standards and Technology (NIST), through a collaboration grant (Number 70NANB12H262). NIST DMLS facilities were used to fabricate the cylindrical test bars.References[1] Anam Md, Pal D., Stucker B., Modeling and Experimental validation of Nickel-based super alloy (Inconel 625) made using Selective Laser Melting. SFF Symposium 2013 proceedings. Page: 463-473.[2] Amato KN, Gaytan SM, Murr LE, Martinez E, Shindo PW, Hernandez J, et al. Microstructures and mechanical behavior of Inconel 718 fabricated by selective laser melting.

In this study, the SLM process was used to fabricate components of TiC/ 316L stainless steel nanocomposite using various laser scanning methods. The results showed that the laser-scanning method used for the SLM process affects the degree of densification, microstructure, and the hardness of the components produced. We believe that the alternative fabrication route presented in this study should significantly increase the use of nanocomposites.Keywords: Selective Laser Melting; Porosity; Microstructure; HardnessIntroduction316L stainless steel alloy is of great interest in the industry for their superior ductility and excellent corrosion resistance [1]. However, the applications of these alloys have been limited due to their poor strength and wear resistance both at room and high temperatures [2, 3].

However, the applications of these alloys have been limited due to their poor strength and wear resistance both at room and high temperatures [2, 3]. To overcome this problem, 316L alloys may be strengthened by a uniform dispersion of very hard fine ceramic particles [4], which can be then used in wear resistant and thermally stable environment.Additive manufacturing (AM) has emerged as a key enabling technology to shorten the product development cycle [6]. AM refers to a class of advanced technologies that can directly use three-dimensional computer aided design data to automatically fabricate multifunctional parts in a layer-by-layer manner [5]. Selective laser melting (SLM) is a new and highly promising AM technique for fabricating complex geometry parts by the selective melting of a metal powder and its subsequent rapid solidification [6].

Selective laser melting (SLM) is a new and highly promising AM technique for fabricating complex geometry parts by the selective melting of a metal powder and its subsequent rapid solidification [6]. The layer-wise addition process and the high thermal-energy gradients of the laser during SLM processing cause the crystals to grow preferentially (i.e., they cause directional solidification) [7]. Therefore, the scanning methods used during the SLM fabrication process can lead to significantly different degrees of densification as well as variations in the microstructure and texture and hence different mechanical properties (i.e., the SLM process can determine the anisotropy of the mechanical properties of the fabricated components).566In a previous work [8], we had studied the effects of the TiC volume content and particle size on the microstructure and mechanical properties of the resulting composite.

In this work, a composite consisting of 316L stainless steel reinforced with 15% (vol.) TiC was produced by SLM using four different laser-scanning methods. The effects of the scanning method used on the solidification microstructure formed were investigated using scanning electron microscopy (SEM), and X-ray diffraction (XRD) analysis. The anisotropy in the mechanical properties of the produced material was analyzed based on the hardness.Experimental proceduresFeedstock powder, and SLM processesThe starting powder materials were: water-atomized 316L having an irregular shape and an average particle size of 45 µm (Fig. 1a), and pure titanium carbide (TiC) with a mean particle size of 1 µm (Fig. 1b). A powder mixture containing 85 vol.% 316L, 15 vol.% TiC was mechanically alloyed under a vacuum atmosphere using a Pulverisette 6 planetary high-energy ball mill. The milling conditions and parameters used for powder preparation are discussed in our previous work [8].

The milling conditions and parameters used for powder preparation are discussed in our previous work [8]. Cylindrical specimens with dimensions of 8 mm x 10 mm were produced according to the SLM processing parameters highlighted in Table 1 and with different scanning methods, shown in Fig. 2.abFigure 1: SEM micrograph of (a) 316L stainless steel powder; (b) TiC powder.Table 1: SLM processing parametersLaser power Atmosphere ScanningLayer thicknessspeed 250 mm/sec 50µm100 WArgon567Figure 2: Scanning methods implemented in the current study.Microstructural and mechanical characterizationsSamples were mechanically grounded and polished with progressively finer grit papers using the standard metallographic techniques and etched in Marble's Reagent for 10 s. The cross- sectional microstructures of the SLM-processed nanocomposites were observed using a Nova 230 Nano scanning electron microscope (SEM).

The cross- sectional microstructures of the SLM-processed nanocomposites were observed using a Nova 230 Nano scanning electron microscope (SEM). XRD measurements were conducted on a PANalytical X'Pert PRO X-ray powder diffractometer (XRD) instrument, operating with a cobalt anticathode at 35 kV and 40 mA using a continuous scan mode at 5º /min. The relative densities were measured by the conventional “Archimedes' principle”.Vickers indentations was performed (Leco, LM800AT) at a load of 200 g at 15 random locations.Results and DiscussionMicrostructural characterizationsFigure 3 shows low-magnification image of the side-view microstructure (i.e. along the building direction) of the SLM-processed nanocomposite using scanning-3 method after etching. The individual long-scan tracks (see the arrows in Fig. 3) and the cross-sections of the melt pools, represented by the half-cylinders/crescents, can be seen clearly. From Fig.

The individual long-scan tracks (see the arrows in Fig. 3) and the cross-sections of the melt pools, represented by the half-cylinders/crescents, can be seen clearly. From Fig. 3, it can be seen that the width of the individual scan tracks/grain width was approximately equal to the hatch spacing (~120 µm.). The widths of the melt pools varied from 110±10 to 170±15 µm, based on the cross-sections of the melt pools; these values were higher than the layer thickness. Further, the melt pools were not always continuous, owing to variations in the depth and shape of the pools.568Figure 3: Cross-sectional microstructure of the SLM-processed TiC/316L nanocomposite processed using scanning-3 method.Figure 4 shows high-magnification side-view (i.e. along the building direction) SEM images of the SLM-processed nanocomposites. In general, all composites sample shows very fine grain size as a result of fast cooling.

along the building direction) SEM images of the SLM-processed nanocomposites. In general, all composites sample shows very fine grain size as a result of fast cooling. The grains are of different sizes, may be due to the varying thermal heat fluxes during every scanning tracks. Apart, more importantly, one can observe the influence of the scanning method used on the resulting microstructure. The grain structures were clearly affected by the geometry of the melting pool, as heat flow and grain growth occurred perpendicular to the liquidus/solidus interface, leading to the morphology observed.In general, the microstructure of SLM-fabricated components is characterized by discontinuous columnar grains/dendrites, whose orientation is determined by the heat-flow direction, which depends on the movement of the laser beam. It can be seen from Fig. 4 that the solidification morphology was highly correlated with the SLM building method.

It can be seen from Fig. 4 that the solidification morphology was highly correlated with the SLM building method. The dendrites had a distinct columnar structure aligned along the building direction, in a manner similar to the case in a directionally solidified microstructure. Further, on comparing Figs. 4a and 4b, it can be seen that the size of the dendrites was significantly affected by laser remelting during double scanning. Refinement in columnar structure is done by laser re-melting which is reported in earlier reports [9-11].With double-pass scanning, the remelting of the solidified layer in the second pass of the laser beam causes the dendrites to become finer than those seen in Fig. 4a. Further, during the second pass, the laser beam is focused on the already-solidified material whereas it is focused on powder particles in the first scan. Thus, in the second pass, less energy is absorbed and heat is conducted away more quickly.

Thus, in the second pass, less energy is absorbed and heat is conducted away more quickly. As a result, the size of the melt pools in the second pass is smaller and some of the material created in the first scan still remains in the final product. A higher number of particle interfaces are broken down in the second pass, and a large number of pore defects are eliminated. The presence of pores retards recrystallization by pinning the dislocations and boundaries necessary for the process [12], and the reduction in porosity during the second pass enhances the recrystallization of the double-pass samples.

The smaller569grain sizes of the double-pass nanocomposite components could be attributed to all of these factors.The samples formed using unidirectional hatches with double scanning (scanning method 3) exhibited two solidification modes (solidification structure consisting of cellular dendrites mixed with equiaxed grains), while in the samples produced using the cross-hatches scanning method (scanning method 4), the solidification microstructure mainly consisted of fine equiaxed grains with a size smaller than 2 µm (Fig. 4d). Further, it was evident that the TiC nanoparticles were present both within and at the grain boundaries.abcdFigure 4: Etched top-view microstructure of the SLM-processed nanocomposite: (a) scanning 1; (b) scanning 2; (c) scanning 3; (d) scanning 4.Table 2 shows the effect of applied scanning methods on the densification level.

It is evident that the building orientation had a pronounced effect on the density, which is typical for SLM processing. Remelting (by performing laser scanning twice) caused more uniform and smoother layers to appear and efficiently reduced the number of pores formed between the neighboring melt pools at the edges of the scan track. When the dual-pass scanning method was used, the voids between the adjacent laser tracks largely disappeared and higher-density parts with a smooth and defect-free surface were produced (Table 2). Previous studies have shown the potential effect of laser remelting on the densification level and surface roughness [10, 11]. It was seen that, with double scanning, the flowability of the melt as well as the rheological properties were enhanced.

It was seen that, with double scanning, the flowability of the melt as well as the rheological properties were enhanced. Hence, the molten 316L with the reinforcing TiC particles are sufficiently high, leading to more homogeneous and efficient mass and heat transfer in the pool.570Table 2: Theoretical relative density of the SLM-processed nanocomposite.Scanning type Relative density (%)1 92.482 96.043 86.914 96.40XRD analysisFigure 5 shows the typical XRD histogram for the SLM-processed nanocomposites using different scanning methods. The XRD pattern of the as-built samples exhibits a significant peak broadening due to the grain refinements as a result of fast cooling. Strong diffraction peaks related to austenitic stainless steel and corresponding to γ-Fe as well as diffraction peaks attributable to the TiC phase were seen in the case of all the samples.

Strong diffraction peaks related to austenitic stainless steel and corresponding to γ-Fe as well as diffraction peaks attributable to the TiC phase were seen in the case of all the samples. This confirmed the formation of TiC-reinforced 316L nanocomposites after SLM processing using different scanning methods. The high cooling rate during the SLM process caused the fabricated nanocomposites to contain an α-Fe phase in a small fraction.Figure 5: XRD histogram for the as-built nanocomposite using different scanning methods.571Mechanical behaviorTable 3 demonstrates the effect of the applied SLM scanning method on microhardness. The observed increase in the hardness could primarily be attributed to grain boundary strengthening and grain refinement and to the presence of the (α-Fe) phase. In general, the scanning methods seemed to have an effect on the hardness values.

In general, the scanning methods seemed to have an effect on the hardness values. Indeed, the porosity level, grain refinement (Hall-Petch relation), and homogeneity of particle distribution (greater homogeneity resulted in greater dispersion strengthening) played significant roles in determining the mechanical properties of the SLM products.

Further, it was evident that the double-pass scanning method resulted in higher hardness; this was because of the additional grain refinement of the fabricated nanocomposite and an improvement in the homogeneity of the particle distribution (i.e., to dispersion hardening being effective) as well as an increase in the density.Table 3: Variation of microhardness as a function of the applied scanning method.Scanning type Microhardness (HV0.2)1 383.872 430.923 370.864 365.3ConclusionsTiC/316L nanocomposites have been successfully produced by SLM process, and the effects of scanning methods on the microstructure and mechanical properties were studied.

The findings of this study would be a valuable reference when selecting the proper scanning method for the SLM fabrication of 316L nanocomposite samples with the desirable grain structure and mechanical properties.AcknowledgmentOne of the authors, Bandar AlMangour, would like to extend his appreciation to the Saudi Basic Industries Corporation, which generously awarded financial support to him.References[1] AL-Mangour B. Powder Metallurgy of Stainless Steel: State-of-the Art, Challenges, and Development. Stainless Steel: Microstructure, Mechanical Properties and Methods of Application: Nova Science Publishers; 2015. p. pp. 37-80. [2] Krishna DSR, Sun Y. Effect of thermal oxidation conditions on tribological behaviour of titanium films on 316L stainless steel. Surface and Coatings Technology 2005;198:447-53. [3] Majumdar JD, Kumar A, Li L. Direct laser cladding of SiC dispersed AISI 316L stainless steel. Tribology International 2009;42:750-3.

FECT OF SECOND-PHASE DOPING ON LASER DEPOSITED Al2O3 CERAMICSF.Y. Niu1, D.J. Wu1, G.Y. Ma1, S.Y. Zhou1, B. Zhang12*1. Key Laboratory for Precision and Non-Traditional Machining Technology of Ministry of Education, Dalian University of Technology, Dalian, Liaoning, China, 1160242. Department of Mechanical Engineering, University of Connecticut, Storrs, CT 06269, USAAbstractDirect fabrication of engineering ceramic components by additive manufacturing (AM) is a relatively new method for producing complex mechanical structures. This study investigates how a second-phase doping may affect Al2O3 ceramic parts deposited by AM with a laser engineered net shaping (LENS) system. In this study, ZrO2 and Y2O3 powders are respectively doped into Al2O3 powders at the eutectic ratio as second-phases to improve the quality of a deposited part.

In this study, ZrO2 and Y2O3 powders are respectively doped into Al2O3 powders at the eutectic ratio as second-phases to improve the quality of a deposited part. The deposited Al2O3, Al2O3/ZrO2 and Al2O3/YAG (yttrium aluminum garnet) parts are examined for their micro-structures and micro-hardness, as well as defects. The results show that doping of ZrO2 or Y2O3 as a second-phase performs a significant role in suppressing cracks and in refining grains of the laser deposited parts. The micro-hardness investigation reveals that the second-phase doping does not result in much hardness reduction in Al2O3 and the two eutectic ceramics are both harder than 1500 Hv. The study concludes that the second-phase doping is good for improving laser deposited ceramic parts.1. IntroductionEngineering ceramics is an irreplaceable material used in the extreme working conditions due to their excellent mechanical properties, as well as thermal and chemical resistances [1].

IntroductionEngineering ceramics is an irreplaceable material used in the extreme working conditions due to their excellent mechanical properties, as well as thermal and chemical resistances [1]. However, traditional manufacturing methods for engineering ceramic parts have difficulties in meeting the requirements from the fast-developing modern technology because of their limitations, such as low sintering efficiency, limited complexity of fabricated structures and relatively low strength [2, 3]. The limitations mostly arise from batching, forming, sintering and machining, which are the main processing steps in the traditional manufacturing methods. Therefore, it is significant to develop new processing methods for making full use of engineering ceramics.Ceramic additive manufacturing is a newly developed method in recent decades and has been successfully demonstrated for its advantages with the use of “indirect” and “direct” methods.

By the indirect method, a green body is first created from ceramic powders or slurry with a high content of an organic or inorganic binder, and then sintered and densified to eliminate binders. Because a high energy source is not needed, many indirect processing methods have been568developed, such as stereo lithography (SL), laminated object manufacturing (LOM), fused deposition of ceramics (FDC) and freeze-form extrusion fabrication (FEF), etc. [4-8]. Through these methods, complex geometries can be obtained, but limitations, such as impurities, porosities and shrinkage, still exist because the processing steps are the same as in the traditional methods. On the other hand, with the direct method, pure ceramic powders can be melted directly by a high energy source and then the melt pool solidifies to form a part. Due to the melting/solidification processes, fully densified ceramic parts with good performances are produced more easily and rapidly.

Due to the melting/solidification processes, fully densified ceramic parts with good performances are produced more easily and rapidly. Comparatively, few direct processing methods, e.g., laser engineered net shaping (LENS) and selective laser melting (SLM), have been developed because of the need for a high energy source [9-14]. Wilkes et al. manufactured ceramic parts from ZrO2/Al2O3 powders with an SLM system and obtained crack-free specimens with flexural strength higher than 500 MPa [9, 10]. Balla et al. used an LENS system to fabricate dense and net-shaped structures of Al2O3 and obtained cylindrical, cubic and gear-shaped parts which showed microstructural anisotropy with hardness of 1550 Hv [11, 12]. Bertrand et al.

Bertrand et al. applied SLS/M to manufacturing net shaped parts from the pure yttria-zirconia powders and demonstrated possibility of processing pure ceramic powders by SLM without doping [13,14].Although the direct methods have demonstrated many advantages in fabricating ceramic parts, there are also technical challenges, such as crack control and property improvements in the laser-aided processes due to the hard and brittle properties of a ceramic material as well as large thermal gradients generated during laser radiation. Preheating was proposed by Wilkes and has been proven to be an effective method for suppressing cracks generated by the SLM fabrication [10]. However, this method provided a limited preheating area and poor surface quality.

However, this method provided a limited preheating area and poor surface quality. Because the direct additive manufacturing methods of ceramic parts have been developed in a relatively short time, they need to be further developed for industrial applications.In this study, an LENS system is used to fabricate ceramic single-bead walls directly from ceramic powders. A new method of second-phase doping is proposed to improve quality of a deposited part. ZrO2 and Y2O3 powders are respectively doped into Al2O3 powders at their respective eutectic ratios as a second-phase to verify the proposed method. Micro-structure and micro-hardness of the fabricated parts are then investigated.2. Experimental proceduresThe LENS experimental system used in this study is shown in Figure 1. The system consisted of a 1000 W Nd:YAG laser, an automatic powder feeder with three containers, and a numerically controlled worktable.

The system consisted of a 1000 W Nd:YAG laser, an automatic powder feeder with three containers, and a numerically controlled worktable. A Ti-6Al-4V plate of 6 mm thick was used as a substrate for ceramic deposition due to its high laser absorptivity and good compatibility with the ceramics. For an efficient use of the ceramic powders, the focus of the powder stream was set on the surface of the substrate. Pure argon was used as a protective gas for separating the fabricated part from the atmosphere.569Figure 1 Experimental system of LENSAl2O3, ZrO2 and Y2O3 ceramic powders (purity wt%>99.5%) of particle sizes in the range of 42-90 m were used for all the experiments after drying at 100 ℃ for 4 hours to eliminate moisture. The three ceramic powders were loaded into the respective containers of a powder feeder. Flow rate of the powders was independently controlled.

The three ceramic powders were loaded into the respective containers of a powder feeder. Flow rate of the powders was independently controlled. In this study, fabrication of Al2O3/ZrO2 parts was conducted by mixing ceramic powders at the eutectic ratio of 58.5 wt% Al2O3 and 41.5 wt% ZrO2 [15]. Similarly, the mixed powders with a ratio of 66.5 wt% Al2O3 to 33.5 wt% Y2O3 were used for fabricating Al2O3/YAG parts [16]. The micrographs of the three powders are shown in Figure 2.During the deposition experiment, the coupled nozzle and laser assembly moved back and forth in the x-y plane to perform ceramic deposition. After each layer deposition, the substrate together with the already deposited part was moved away from the nozzle by one layer thickness along the z direction so as to maintain a constant focus position for depositing the next layer. The process was then repeated to finish the single-bead wall. For different materials, the process parameters were different.

The process was then repeated to finish the single-bead wall. For different materials, the process parameters were different. Based on the systematic experiment, the optimum process parameters were obtained for the three powder materials, as shown in Table 1.Figure 2 Ceramic powders used in the experiment (a) Al2O3; (b) Y2O3; (c) ZrO2570Table 1 Process Parameters for the Ceramic PowdersMaterialsAl2O3Laser Power (W) 350Nozzle Travel Speed (mm/min) 300Powder Flow Rate (g/min) 1.36Z-increment (mm) 0.22Al2O3/Y2O33203501.08/0.550.18Al2O3/ZrO24104001.22/0.870.25The fabricated ceramic parts were cut from the longitudinal cross-section A-A, as shown in Figure 3, and then prepared by coating a thin Au to observe the microstructure in a Scanning Electron Microscope (SEM). For the Al2O3/ZrO2 and Al2O3/YAG ceramic parts, their cross-sections were polished with a diamond disk and an abrasive paper before being coated with Au.

For the Al2O3/ZrO2 and Al2O3/YAG ceramic parts, their cross-sections were polished with a diamond disk and an abrasive paper before being coated with Au. Vickers micro hardness measurements were made on the polished samples using a 1000 g load for 15 seconds, and an average value of 10 measurements on each sample was reported.Figure 3 The longitudinal cross-section of a fabricated part3. Results and discussion3.1 Properties of deposited Al2O3 partFigure 4(a) is a photographic view of a 50-layer Al2O3 single-bead wall fabricated from the Al2O3 powders under the process parameters shown in Table 1. Many stripes could be observed on the surface of the part, and cracks were found along the deposition direction, most of which in the middle portion of the part. On the other hand, a few cracks were found near the two side edges of the single-bead wall, but much longer than those in the middle portion.

On the other hand, a few cracks were found near the two side edges of the single-bead wall, but much longer than those in the middle portion. It should be noted that very few cracks were generated along the scanning direction.Crack generation in Figure 4(a) may mainly be related to two factors: thermal stress during the laser depositing process and crystallographic orientation of the fabricated part [17]. Because laser scanning is a layer-by-layer process, highly non-uniform temperature distribution is formed across the deposited part. During the solidification process, volume contraction in the laser571scanning direction is constrained by the previously deposited layer with the first layer being constrained by the substrate. Consequently, thermal tensile stress occurs in the part along the laser scanning direction.

Consequently, thermal tensile stress occurs in the part along the laser scanning direction. Substrate temperature is usually lower than that of the ceramic part due to the good thermal conductivity of the substrate, resulting in the tensile stress in the first layer, which in turn generates a bending moment, as shown in Fig. 5. The thermal stress distribution of the fabricated ceramic part is similar to that of the metal parts, which has been reported in the literature [18-19]. On the other hand, it can be seen from Figure 4(b) that the microstructure of the fabricated part was comprised of directional columnar crystals along the deposition direction with an intergranular space of about 10-15 m due to the directional heat dissipation [20]. It is easy for cracks to propagate along the crystal boundaries in the deposition direction, whereas it may be much harder to propagate along the scanning direction because more energy is needed for crack propagation through a crystal.

Consequently, it is more preferable for a crack to form and propagate in the vertical direction than in the other directions. For the above reasons, vertical cracks are preferably generated (Figure 4(a)).Figure 4 (a) Vertical cracks and (b) Microstructure of the deposited Al2O3 partFigure 5 Bending moment and tensile stress induced by solidification contractionThe average micro-hardness of the deposited Al2O3 part was measured 1800 Hv, compared to 1600 Hv for the traditional Al2O3 ceramic materials. The hardness essentially rests with the572corundum structure of α-Al2O3 and its ionic (63%) and covalent (37%) bonding energies [21], whereas the LENS method results in the dense structure of a deposited part, which is because of its unique melting-solidification process.3.2 Properties of deposited Al2O3 /YAG partFigure 6(a) presents a photograph of a 50-layer Al2O3/YAG single-bead wall fabricated from the mixed ceramic powders at the eutectic ratio.

During the fabrication process, the following chemical reactions could be used to describe the formation of Y3Al5O12 (YAG) [22]. Then the rest of Al2O3 and the generated YAG should solidify simultaneously at 1826 ℃ to form the eutectic ceramics of Al2O3/YAG [22].Al2O3 + 2Y2O3 → Y4Al2O9(YAM) (900 − 1100 ℃) (1) Y4Al2O9 + Al2O3 → 4YAlO3(YAP) (1100 − 1250 ℃) (2) 3YAlO3 + Al2O3 → Y3Al5O12(YAG) (1400 − 1600 ℃) (3)Compared with the Al2O3 part discussed above, a two phase ceramic part was deposited with fewer cracks. There were three vertical cracks in the depositing direction and two horizontal cracks in the scanning direction, which might be induced by a tensile stress in the corresponding directions. Shown in Figure 6(b), light-grey YAG phase evenly embeded in the Al2O3/YAG eutectic matrix.

Shown in Figure 6(b), light-grey YAG phase evenly embeded in the Al2O3/YAG eutectic matrix. The volume fraction of the two phases did not seem to match the material ratio at the eutectic point, which might be due to the local evaporation of Al2O3 or the rapid solodification during the deposition process. The microstructure of the Al2O3/YAG single-bead wall was finer without obvious directional growth characteristics, which could help surpress both vertical and horizontal cracks. From the magnified image of Figure 6(b), a fine-grained microstructure with a eutectic spacing smaller than 1 m is observed, resulting in the extremely complex interphase boundaries.

From the magnified image of Figure 6(b), a fine-grained microstructure with a eutectic spacing smaller than 1 m is observed, resulting in the extremely complex interphase boundaries. Therefore, it would be difficult for a crack to initiate and propagate in such a fine-grained microstructure.Figure 6 (a) Vertical and horizontal cracks and (b) Microstructure of Al2O3/YAG part573Based on the experimental measurement, the average micro-hardness of Al2O3/YAG was 1575 Hv, slightly lower than that of the Al2O3 ceramic part. The value was comparable to that of the traditional Al2O3/YAG eutectic ceramics made by the directional solidification or the laser-heated floating zone method [23,24].3.3 Properties of deposited Al2O3 /ZrO2 partFigure 7 (a) Fabricated Al2O3/ZrO2 part without cracks; (b) Microstructure of the Al2O3/ZrO2 partFigure 7(a) provides a photograph of a 60-layer Al2O3/ZrO2 single-bead wall.

The eutectic Al2O3/ZrO2 part was successfully deposited with no cracks along both the depositing and the scanning directions. Compared with the deposited Al2O3/YAG ceramic part, the microstructure of Al2O3/ZrO2 was also consisted of two phases, the black Al2O3 phase and the white ZrO2 phase, but the eutectic spacing of Al2O3/ZrO2 was nano-sized which was obviously finer than that of the eutectic Al2O3/YAG part in Figure 7(b). As a result, the eutectic Al2O3/ZrO2 part can be made without cracks.The micro-hardness of the traditional ZrO2 is approximately 1200 Hv.

As a result, the eutectic Al2O3/ZrO2 part can be made without cracks.The micro-hardness of the traditional ZrO2 is approximately 1200 Hv. However, the average hardness of the deposited eutectic Al2O3/ZrO2 part reached 1715 Hv, although lower than that of the pure Al2O3, but higher than that of the Al2O3/ZrO2 parts (an average value of 1460 Hv) made by the traditional methods.3.4 Parts deposited for demonstration purposeAs mentioned above, the second-phase doping has been proven to be an effective method for suppressing cracks and refining microstructures during the direct fabrication process of ceramics by LENS. For the demonstration purpose, more eutectic Al2O3/ZrO2 parts were deposited by this method. As shown in Figure 8, a cylindrical structure longer than 50 mm, an arc wall and a short cylinder were successfully fabricated with no cracks but good surface quality.574Figure 8 Al2O3/ZrO2 parts of (a) Arc wall; (b) Long cylinder; (c) Short cylinder4.

SummaryDirect fabrication of Al2O3 ceramic single-bead walls is conducted by the LENS system in this study. Y2O3 and ZrO2 are doped into Al2O3 powders respectively as a second-phase to restrict crack formation and improve microstructure of the deposited ceramic parts. The results show that both the second-phase doping of Y2O3 and ZrO2 perform well in refining the microstructures of the direct laser deposited Al2O3/YAG and Al2O3/ZrO2 ceramic parts, which suppresses crack formation. The micro-hardness investigation reveals that the second-phase doping does not result in much hardness reduction in Al2O3 and the two eutectic ceramics are both harder than 1500 Hv.

The micro-hardness investigation reveals that the second-phase doping does not result in much hardness reduction in Al2O3 and the two eutectic ceramics are both harder than 1500 Hv. The study indicates that second-phase doping is an effective way to improving laser deposition quality of ceramic parts.AcknowledgementsThe authors would like to acknowledge the financial support from the National Natural Science Foundation of China (No.51175061), the Science Fund for Creative Research Groups (No.51321004) the China Postdoctoral Science Foundation Funded Project and (No.2014M551072).575References[1] D. Richerson, “Modern Ceramic Engineering: Properties, Processing, and Use in Design,”CRC press, 2005.[2] W. D. Kingery,, H. K. Bowen, D. R. Uhlmann, Introduction to Ceramics, 2nd Ed., 1976 [3] M. N. Rahaman. Ceramic Processing. John Wiley & Sons, Inc., 2006. [4] H. C.

fect of section thickness and build orientation on tensile properties and material characteristics of Laser Sintered nylon-12 partsMajewski, C.E. and Hopkinson, N.Additive Manufacturing Research Group, Loughborough University, Loughborough, United Kingdom, LE11 3TUAbstractIt has been suggested that different section thicknesses in Laser Sintered parts may cause variations in mechanical properties, and this has previously been demonstrated for some properties (e.g. fracture toughness). The research presented here investigates whether the same is true of tensile properties, and whether the orientation of parts within the build volume has any effect on this.

fracture toughness). The research presented here investigates whether the same is true of tensile properties, and whether the orientation of parts within the build volume has any effect on this. Results are presented for three different orientations of tensile specimens, at a range of thicknesses from 2mm to 6mm, showing that, at any of the orientations tested, the section thickness had no significant effect on any of the main tensile properties.These results are in direct contradiction with related research investigating the effect of section thickness on Fracture Toughness, where an increase in thickness also increased the toughness of the parts.

This highlights the importance to Additive Manufacturing users of identifying the correct properties to assess when choosing a suitable process or material, and when designing complex parts.IntroductionContextWhilst Additive Manufacturing (AM) technologies have to an extent become accepted by industry, their uptake has been slower than could be expected, largely due to a lack of knowledge regarding how to design for the relevant processes, and how components will behave in practice. Datasheet values of properties go some way to improving this situation, but are often regarded as providing insufficient information for ‘real life’ use.

In addition, a substantial amount of anecdotal evidence is common within the AM community, much of which is either not fully accessible, or has not yet been proven.In order for current and potential users of AM to take full advantage of the many benefits of these technologies, more information must be obtained, and published, regarding the use, and expected behaviour, of parts produced using various AM technologies.The research presented here was performed as part of the £350,000 ‘Advanced Understanding and Control of Polymer Sintering’ (AUCPS) project, funded by Loughborough University’s Innovative Manufacturing and Construction Research Centre (IMCRC)1, and in collaboration with Electro Optical Systems (EOS GmbH)2, Burton Snowboards3, and The University of Louisville4.Whilst many acknowledge the wide range of benefits offered by AM technologies, and in particular Laser Sintering (LS), and while there has been a considerable amount of published literature in these areas, many of the nuances of the LS process are still not fully understood.The overall aim of the project is to achieve a greater understanding of the factors affecting the properties of polymer Laser Sintered parts, and in turn to use this understanding to control the behaviour of these parts, and to provide AM users and designers with the knowledge to achieve suitable performance from parts and components.Effect of geometrySubstantial amounts of research have been carried out into benchmarking of geometric capabilities for most AM systems, in terms of feature size resolution, surface finish and accuracy5,6.

Additionally, a large amount of published data can be found relating to the effects of varying processing parameters on the mechanical properties that can be achieved7,8,9.However, little investigation has been made into the effects of geometry on the mechanical properties of the parts themselves. Datasheet values for many commonly required properties are quoted by many manufacturers, but only for standard specimen sizes. It has been suggested that, due to different thermal conditions when building parts of varying section thicknesses, the achievable properties may be markedly different. In order to gain increased acceptance, and uptake, of Laser Sintering, it is crucial to improve knowledge of the differences, if any, to be expected throughout a component featuring different cross-sectional areas.Related research – Fracture ToughnessPrevious research within the AUCPS project has investigated the effect of section thickness on Fracture Toughness10.

Laser Sintered single edge notch bending (SENB) beam specimens were produced at a range of thicknesses (2 mm to 10 mm), and showed an increase (~48 %) in the energy required to initiate crack growth per unit area with increasing thickness for LS Nylon-12 samples.Laser Sintered samples produced at a range of thicknesses showed a small increase (~5 %) in density when increasing thickness of specimens from 2 mm to 10 mm, possibly due to improved heat retention in the thicker specimens, leading to an increase in the effective melt temperature, and subsequently to improved melt flow, and therefore to more effective sintering.

A slight increase in crystallinity was also observed as section thickness increased.However, these small increases in density and crystallinity were not considered high enough to cause the large increase in toughness observed, and it was suggested that increases in molecular weight in thicker specimens, due to increased time at elevated temperatures, could be the reason for these differences.MethodologyTo complement the Fracture Toughness investigations described in the previous section, and as tensile properties are among the most commonly quoted for LS Nylon-12 parts, the following package of work was designed to assess the effects of section thickness on tensile properties. Tensile specimens were produced and tested at a range of thicknesses.

Three different orientations were also assessed, in order to establish whether this in any way accentuated any effects of section thickness.Part geometryTensile test specimens were designed in accordance with BS EN ISO 527-2:199611, as shown in Figure 1.Figure 1 - Tensile test specimen (all dimensions in mm)Build layoutTensile test specimens were produced at six different nominal thicknesses (2mm, 3mm, 4mm, 5mm and 6mm) and in three different orientations – YX, YZ and ZY (where the first letter denotes the axis parallel to the longest dimension, and the second letter denotes the axis parallel to the second longest dimension).Six parts were produced at each thickness and each orientation, five of these for tensile testing, and a further sample of each for any additional material testing required at a later date.Figure 2 shows the layout of parts within the build volume.Figure 2 - Build LayoutLaser Sintering parametersParts were produced in PA2200 (a Nylon-12 based polymer provided by EOS), on an EOS Formiga P100 system.

All parts were produced using the manufacturer’s default processing parameters for this material, as shown in Table 1.ParameterSetpointLayer thickness Part bed temperature Removal chamber temperature Laser power (contour) Scan speed (contour) Laser power (hatching) Scan speed (hatching) Scan spacing Pre-heat time0.1mm 172 °C 159 °C 16 W 1500 mms-1 21W 2500 mms-1 0.25 mm 2 hoursTable 1 - LS processing parametersOnce building was complete the parts were left to cool to 50°C before removing from the part cake.Tensile testingAll specimens were conditioned at 20 °C (+/−1 °C) and 50% (+/−5%) relative humidity prior to testing. Tensile tests were performed on the specimens using a Zwick Z030 tensile testing machine fitted with a long travel contact extensometer. The Young’s Modulus was measured using a 1mm/min strain rate to 0.25 %, following which the Tensile Strength and Elongation at Break were measured at 5mm/min until failure of the specimen.

The Young’s Modulus was measured using a 1mm/min strain rate to 0.25 %, following which the Tensile Strength and Elongation at Break were measured at 5mm/min until failure of the specimen. The Zwick software, TestXpert, was used to enable the calculation of the Tensile Strength, Young's Modulus and Elongation at Break of each of the parts.ResultsSpecimen thicknessDuring testing it was noted that in every case the actual measured thickness of the specimens was somewhat larger than the nominal value, suggesting that the scaling factors set within the P100 operating software require adjustment.

This was true to the greatest extent when producing parts in the YX orientation, whereby the actual thickness was on average 9.8% larger than the nominal value, as compared with an average of 3.5% larger in YZ and ZY.However, as the calculated values for tensile properties, presented in the following sections, take into account the actual thickness of the specimen, these differences between actual and nominal dimensions will not affect the results shown.Tensile StrengthFigure 3 shows the Tensile Strength for each of the thicknesses and orientations tested.Figure 3 - Tensile Strength at different thicknesses and orientationsIt can be seen from Figure 3 that, within the range tested, there is no obvious effect of section thickness on Tensile Strength at any of the orientations tested.In order to further assess any relationship between the section thickness and Tensile Strength, the Coefficient of Determination (R2) value was calculated for each orientation (see Table 2 for values).

A linear regression analysis indicated a slight increase in Tensile Strength with increasing thickness, but it can be seen that all values of R2 fall well below the value of 0.9 generally accepted as indicating a strong correlation, confirming that there is no significant effect of section thickness on Tensile Strength.Orientation YX YZ ZYR2 0.00 0.06 0.65Table 2 - R2 values for correlation between thickness and Tensile StrengthFigure 4 shows the effect of orientation on Tensile Strength. All specimens produced in each orientation were averaged and included here.Figure 4 - Effect of orientation on Tensile StrengthThe results shown in Figure 4 demonstrate that the orientation of parts had no apparent effect on the Tensile Strength.

The large range in results for the YX and YZ conditions can be explained by two anomalous values for each condition, and therefore is considered random rather than of any significance.Young’s ModulusFigure 5 shows the Young’s Modulus for each of the thicknesses and orientations tested.Figure 5 - Young's Modulus at different thicknesses and orientationsThe results show in Figure 5 indicate that, again within the range tested, there is no obvious effect of section thickness on Young’s Modulus at any of the orientations tested.Table 3 shows the R2 values generated for each orientation, and it can be seen that all values of were less than 0.1, further confirming the lack of any relationship between section thickness and Young’s Modulus.Orientation YX YZ ZYR2 0.02 0.08 0.02Table 3 - R2 values for correlation between thickness and Young's ModulusFigure 6 shows the effect of orientation on Young’s Modulus.Figure 6 - Effect of orientation on Young's ModulusIt can be seen from Figure 6 that, as with Tensile Strength, part orientation had no significant influence on the overall values recorded.

Whilst the values recorded for the YZ direction show a small increase in the average value, the wide range of results for each condition, and the large overlap in range, means that no clear statement of effect can be made.Elongation at BreakFigure 7 shows the Elongation at Break for each of the thicknesses and orientations tested.Figure 7 - Elongation at Break at different thicknesses and orientationsAs with the Tensile Strength and Young’s Modulus, it can be seen that there is no apparent influence of section thickness on Elongation at Break.Table 4 shows the R2 values obtained for each orientation.

It can be seen that for every orientation the R2 value is 0.10 or less, once again confirming that there is no significant effect of section thickness on this property.Orientation YX YZ ZYR2 0.10 0.09 0.03Table 4 - R2 values for correlation between thickness and Elongation at BreakFigure 8 shows the effect of orientation on Elongation at Break.Figure 8 - Effect of orientation on Elongation at BreakIn this case, as would be expected based on previous research12, the samples produced in the ZY orientation showed a lower average Elongation at Break (~ 0.65 that of the YX and YX orientations). The range of results achieved was relatively constant across all three conditions.Discussion and ConclusionsWhen considering Tensile Strength and Young’s Modulus, the orientation of the parts showed no demonstrable effect on the values achieved, demonstrating that they are more robust to changes in build layout.

The same was true for the Elongation at Break in the YX and YZ direction, with a decrease in the property when parts were produced in the ZY direction. Results also show that, within the range tested, section thickness has no significant effect on tensile properties at any of the orientations tested, in contradiction with the effects previously identified for Fracture Toughness.This would indicate that the small increases in density and crystallinity previously reported at larger section thicknesses are not large enough to have a significant effect on tensile props. Although not yet confirmed, the expected increase in molecular weight caused by the higher time at elevated temperature for thicker cross-sections, appears to also have no significant effect on tensile properties.

These results indicate that the Fracture Toughness is substantially more susceptible to variations in these properties than the Tensile Properties of LS parts.Additional research within the AUCPS project13 has since demonstrated a decrease in molecular weight between LS parts produced early in a build and those produced closer to the end of the same build. This clearly adds weight to the argument that increased time at elevated temperature causes an increase in the molecular weight of parts. Further work is planned to investigate thesuggestion that thicker parts do indeed possess higher molecular weight, and also that it is this difference in molecular weight that causes the variations in Fracture Toughness.

The inclusion of additional tensile test specimens within this work will also enable confirmation (or otherwise) that the tensile properties are substantially less susceptible to changes in molecular weight than the Fracture Toughness.In terms of practical implications of this research, the results presented indicate that the tensile properties of LS Nylon-12 parts are particularly robust to changes in geometry and or orientation, providing AM users with a high level of freedom over their designs.

The one exception is when considering Elongation at Break, where care must be taken to avoid orientating in the Z direction if the highest properties are to be achieved.Finally, the research presented here, and the different trends seen for different mechanical properties, has highlighted the importance for AM users of clearly identifying the most crucial properties to identify for their required application.AcknowledgementsThis work was supported by EPSRC/IMCRC project 251.References1 http://www.lboro.ac.uk/eng/research/imcrc/ 2 http://www.eos.info/ 3 http://www.burton.com 4 http://louisville.edu/ 5 Mahesh, M., Wong, Y.S., Fuh, J.Y.H. and Loh, H.T., Benchmarking for comparative evaluation of RP systems and processes, Rapid Prototyping Journal, V10, N2, · 2004, · pp123 – 135 6 Kim, G.D. and Oh, Y.T., A benchmark study on rapid prototyping processes and machines: quantitative comparisons of mechanical properties, accuracy, roughness, speed, and material cost, Proc.

and Hopkinson, N., Fracture mechanics approach to compare Laser Sintered parts and injection mouldings of Nylon-12, paper submitted to Proceedings of the Institution of Mechanical Engineers, Part B, Journal of Engineering Manufacture, June 2010 11 BS EN ISO 527-2:1996, Plastics - Determination of Tensile Properties - Part 2: Test Conditions for Moulding and Extrusion Plastics12 Gibson, I. and Shi, D., Material properties and fabrication parameters in the selective laser sintering process, Rapid Prototyping Journal, V3, N4, 1997, pp 129-136 13 Haworth, B. and Hopkinson, N., Advanced understanding and control of polymer sintering, proceedings of the International Conference on Additive Manufacturing, Loughborough University, UK, July 2010.

FECT OF SEGREGATED FIRST AND SECOND MELT POINT ON LASER SINTERED PART QUALITY AND PROCESSINGBen Fulcher, David K. LeighHarvest Technologies Belton, TX 76513AbstractEfforts to tailor laser sintering polymers to enhance part quality, performance, and processing have relied on the characterization of the polymers using Melt Flow Index (MFI) and Differential Scanning Calorimetry (DSC). Two grades of laser sintering nylon polyamide are compared and the resultant processing window, part quality, and mechanical behavior are discussed. A better understanding of characterization techniques and the processing of laser sintered polymers is leading to engineering thermoplastics for exclusive use in additive manufacturing.1. IntroductionThermal characteristics of a given polymer are essential in determining its suitability for use in Selective Laser Sintering (SLS). Semi-crystalline thermoplastics typically display a discrepancy between melting and recrystallization temperatures.

Semi-crystalline thermoplastics typically display a discrepancy between melting and recrystallization temperatures. For SLS processing, the range of temperatures that exists for a given thermoplastic between the recrystallization and melting temperatures is considered the operating temperature window of the material. Thus, the material should be kept within this temperature range during SLS processing. By keeping the sintered material in a molten state above the recrystallization temperature, part warping due to thermal variation among sintered layers is reduced. By keeping the powder material in a solid state below the melting temperature, undesired melting of surrounding powder is avoided. Once the entire SLS build is sintered, the part bed is cooled and the sintered material recrystallizes and solidifies. Differential Scan Calorimetry, or DSC, can be used to capture the melting and crystallization temperature ranges of polymers.

Differential Scan Calorimetry, or DSC, can be used to capture the melting and crystallization temperature ranges of polymers. A typical DSC curve for a semi-crystalline polymer is displayed in Figure 1, where heat capacity at constant pressure of the material is plotted versus the temperature of the material.A polyamide 11 powder under the brand name Rilsan® Invent Natural (referred to as Invent in this paper) has recently been engineered for use in laser sintering by Arkema Inc. While similar to another polyamide 11 product under the brand name Rilsan® D80 Natural ES (referred to as D80 in this paper), the Invent powder has undergone development to tailor its thermal characteristics to the laser sintering process. Among various improvements over the D80 powder made by Arkema Inc., such as optimized particle size distribution, a key improvement was the introduction of a melting-point transition.

Among various improvements over the D80 powder made by Arkema Inc., such as optimized particle size distribution, a key improvement was the introduction of a melting-point transition. First and second melting points exist in the Invent material, leading to increased ease of processing and enhanced recyclability for SLS. Preliminary results have also shown improved mechanical properties over similar polymers processed in SLS. From a thermal standpoint, a primary difference between the Invent and D80 materials is the occurrence of this melting-point transition in the Invent powder, whereas the D80 powder maintains a relatively constant melting point after a melt and recrystallization cycle.556Once the Invent material undergoes melting at a temperature of roughly 203°C, the melting temperature drops to roughly 188°C. Thus, the material is said to display a melting-point transition from a first melting point (203°C) to a second melting point (188°C).Figure 1.

Thus, the material is said to display a melting-point transition from a first melting point (203°C) to a second melting point (188°C).Figure 1. Typical DSC curve for a semi-crystalline thermoplastic2. Single Melting-Point MaterialsIn SLS processing of materials with a single melting point, the powder bed is kept at a temperature level within the operating range of the material, preferably close to, but below, the melting temperature. Thus, the laser-melted powder remains molten and fluid while the surrounding powder remains solid. However, since the surrounding powder is also kept just below its melting temperature, the process often results in relatively poor feature detail due to solid state sintering of the surrounding powder to the laser-melted material [1]. If the melted material is allowed to approach its crystallization temperature, each melted layer is more likely to experience densification, leading to increased likelihood of part warping.

If the melted material is allowed to approach its crystallization temperature, each melted layer is more likely to experience densification, leading to increased likelihood of part warping. The photograph in Figure 2 shows the effect of this solid-state sintering on the edge of a melted region. At the top of the photograph, solid particles can be seen that are fused to the edge of the part.In addition, the surrounding powder is highly prone to thermal aging, a common issue with SLS materials as it hinders material recyclability. As the powders are cycled from room temperature to near the melting temperature, the fine particles experience solid-state sintering with each other, disrupting the desired particle size distribution of the material [1]. This effect can lead to poor powder flow characteristics as well as poor surface quality known as “orange- peel,” displayed in Figure 3.

This effect can lead to poor powder flow characteristics as well as poor surface quality known as “orange- peel,” displayed in Figure 3. Furthermore, higher temperatures can cause polymer chains to grow, thereby increasing the melting temperature of the used powder [2].557Figure 2. Effect of solid-state sintering on edge of part [3]Orange PeelFigure 3. Orange peel surface texture on SLS part [2]3. Melting-Point TransitionWhile some semi-crystalline thermoplastics can contain multiple melting points due to the inclusion of differing crystal structures [4], melting-point transitions, such as the one displayed by the Invent material, could be the result of thermal pre-processing. For example, one558method of inducing a melting-point transition in a semi-crystalline polymer is through thermal annealing. By annealing the polymer at a constant temperature below, but near, the melting point of the polymer, crystallinity within the material can be increased.

By annealing the polymer at a constant temperature below, but near, the melting point of the polymer, crystallinity within the material can be increased. Driven by thermodynamic forces, amorphous chains are incorporated into the lamellae, thickening the lamellae. Since the melting temperature of lamellae increases with lamellae thickness, the annealing process increases the overall melting temperature of the material [5]. This increased melting temperature is referred to as the first melting point of the material for SLS processing. When the annealed powder is then melted in SLS, the melting temperature of the laser-fused material drops to its original pre-annealed value. The pre-annealed melting temperature is referred to as the second melting point of the material for SLS processing.

The pre-annealed melting temperature is referred to as the second melting point of the material for SLS processing. Thus, the material is pre-processed through thermal annealing, thereby raising the melting temperature, converted into a fine powder, and then selectively melted in the SLS process, returning the melting temperature to its pre-annealed state.The segregation of first and second melting points in the Invent powder provides various benefits for SLS. Unlike typical powders, where the part bed temperature is kept just below the melting temperature of the material, Invent powder can be successfully processed in SLS at temperatures well below its melting point. This is because once the desired material is laser-melted, the melting point of the material transitions to a lower temperature, keeping the molten material from experiencing recrystallization. To further understand this transition, examine the DSC curves displayed in Figure 4, provided by Arkema Inc.

To further understand this transition, examine the DSC curves displayed in Figure 4, provided by Arkema Inc. As seen in the image, the first heating curve shows a melting point of roughly 203°C. The molten material then follows curve 2 for cooling, experiencing recrystallization at roughly 157°C. When the material is heated a second time, it experiences melting at roughly 188°C. Thus, for SLS processing, the laser-melted material remains highly fluid since it can be held above the second melting temperature while still remaining well below the melting temperature of the powder. Therefore, minimal to no part warping due to layer shrinkage is expected.Figure 4. DSC curves for Rilsan® Invent Natural powderIn Figure 4, labels 1, 2 and 3 represent the order in which the processes occur. Note also that exothermic processes are considered positive in terms of heat flow, whereas endothermic559processes are considered negative.

Note also that exothermic processes are considered positive in terms of heat flow, whereas endothermic559processes are considered negative. In contrast to the DSC curves for the Invent powder, the DSC curves for the single melting-point D80 powder are displayed in Figure 5. Although the shape of the DSC curve changes from the first to the second heating, the melting temperature remains very similar. Thus, the D80 material is considered to be a single melting-point material.Figure 5. DSC curves for Rilsan® D80 Natural ES powder4. Preliminary Trials with Rilsan® InventPreliminary trials with Invent powder have shown promising results.

DSC curves for Rilsan® D80 Natural ES powder4. Preliminary Trials with Rilsan® InventPreliminary trials with Invent powder have shown promising results. With the part bed temperature held at 192°C, well below the first melting temperature of 203°C, SLS has produced parts with good mechanical properties, very little warping, good feature detail, and good recyclability.Tensile bars built along the z-axis of the build chamber, using D80 powder, had an average elongation at fracture of 26.7% and an average ultimate tensile strength of 7323 psi. Comparatively, z-oriented tensile bars built using Invent powder had an average elongation at fracture of 43.4% and an average ultimate tensile strength of 7855 psi. The improved mechanical properties displayed by the Invent powder are the result of a high level of layer-to-layer adhesion.

The improved mechanical properties displayed by the Invent powder are the result of a high level of layer-to-layer adhesion. This good adhesion is likely, at least in part, due to the fact that the laser-melted Invent material is held above its melting point, keeping the bars in a highly molten state for an extended period of time. Tensile data from two bars, one from each build, is plotted in Figure 6. The two bars chosen for plotting displayed the highest elongation at fracture for their respective builds.560Figure 6. Tensile testing curves for select Invent and D80 specimensPrototype parts built with the Invent material showed little warping. As expected, since the molten material is held far above its crystallization temperature, layer shrinkage during the build is eliminated. Thus, stresses in the parts due to layer shrinkage, which cause in-build curling of the parts, did not arise in the Invent powder trial.

Thus, stresses in the parts due to layer shrinkage, which cause in-build curling of the parts, did not arise in the Invent powder trial. A prototype created with the Invent material is shown on a flat surface in Figure 7.Figure 7. Prototype built with Rilsan® Invent Natural powder displaying minimal warpingThe Invent powder was also able to produce parts with good feature detail. Small extruded lettering is often difficult to produce in SLS due to a number of reasons, including solid-state sintering of surrounding powder to the borders of the letters. However, because the561Invent powder was held well below its melting temperature throughout the build, solid-state sintering of powder to the edge of the parts did not appear to be as prevalent in the Invent powder as with other SLS materials. An image of small lettering on one of the prototype parts can be seen in Figure 8.

An image of small lettering on one of the prototype parts can be seen in Figure 8. Although somewhat difficult to read in the photograph due to the white letters and white back-drop, the phrase “Low Voltage Circuit Breakers” was clearly produced on the part.Figure 8. Feature detail using Rilsan® Invent Natural powderAlthough a true study on recyclability of SLS powder would involve numerous trials in which the ratio of virgin to used powder is varied, preliminary trials of the Invent powder indicate good recyclability. A qualitative indication of recyclability is the “feel” of the part cake after a build is processed in SLS. If the part cake, which includes all of the powder surrounding the parts, feels soft, it is likely that very little solid-state sintering amongst the powder particles occurred. This typically indicates good recyclability, as the particle size distribution remains relatively unaltered.

This typically indicates good recyclability, as the particle size distribution remains relatively unaltered. However, if the part cake feels hard, the powder particles likely experienced significant solid-state sintering, resulting in an increase of average particle size. This indicates poor recyclability as larger particles reduce dimensional accuracy and alter the thermal characteristics of the powder. Trial builds with Invent powder have had a very soft part cake, indicating good recyclability. The photograph in Figure 9 shows the part cake from an Invent trial.562Figure 9. Part cake using Rilsan® Invent Natural powderIn preliminary trials, the Invent powder has proven relatively easy to process. Because the part cake is soft, the time required to remove parts from the cake, as well as remove excess material by bead blasting and sanding, is reduced. Furthermore, the temperature of the part bed does not need to be tightly controlled to still produce good mechanical properties.

Furthermore, the temperature of the part bed does not need to be tightly controlled to still produce good mechanical properties. However, when the part bed temperature was held at 189°C, curling of parts was noticeable. Further studies should be conducted to fully determine the effect of part bed temperature on mechanical properties and appearance.5. ConclusionsMelting point transitions in semi-crystalline thermoplastics have been successfully utilized in SLS to enhance part quality, mechanical properties, material recyclability, and ease of processing. By harnessing the ability of the recently developed Rilsan® Invent Natural to change melting temperatures after a first melt, preliminary trials have displayed good results in each of these areas. The melting point transition allows molten material to be held above its melting temperature, and non-molten powder to remain well below its melting temperature.

The melting point transition allows molten material to be held above its melting temperature, and non-molten powder to remain well below its melting temperature. These effects reduce the likelihood of in-build curling, improve layer-to-layer adhesion and therefore mechanical properties, and reduce the degree of solid-state sintering among loose powder, thereby improving feature detail and powder recyclability. Further work should be done to collect data on material recyclability and ease of processing.563References[1] J. Roesler, H. Harders, and M. Baeker, Mechanical Behavior of Engineering Materials.Springer, 2007, pp. 106-107.[2] J. Kruth, G. Levy, R. Schindel, T. Craeghs, and E. Yasa, “Consolidation of Polymer Powders by Selective Laser Sintering,” in Proceedings of 3rd International Conference on Polymers and Moulds Innovations, 2008, pp. 15-30.[3] H. Zarringhalam, N. Hopkinson, N. F. Kamperman, and J. J.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceTHE EFFECT OF SHEAR-INDUCED FIBER ALIGNMENT ON VISCOSITY FOR 3D PRINTING OF REINFORCED POLYMERSDylan Hoskins1, Christine Ajinjeru1, Vlastimil Kunc2, John Lindahl2, Daniel Moreno Nieto3, Chad Duty1,21 University of Tennessee, Knoxville 2 Manufacturing Demonstration Facility, Oak Ridge National Laboratory 3 University of Cadiz, SpainAbstractMaterial printed with large scale additive manufacturing systems such as the Big AreaAdditive Manufacturing (BAAM) system experience a wide range of shear rates during the extrusion process. The shear rate can vary over five orders of magnitude as the material passes through the single screw extruder and is deposited onto previous layers.

The shear rate can vary over five orders of magnitude as the material passes through the single screw extruder and is deposited onto previous layers. When fiber reinforced materials are deposited, the fibers can become highly aligned in the direction of flow due to the high shear stresses experienced as the material passes through the nozzle. Therefore, accurate analysis of the viscoelastic response of a polymer during extrusion should replicate these conditions as closely as possible. This study evaluates the effect of a pre-conditioning shear strain on the extrusion viscosity of carbon fiber reinforced acrylonitrile butadiene styrene (ABS).IntroductionDeveloping new materials and identifying existing materials that can be used forextrusion-based additive manufacturing (AM) processes, such as fused filament fabrication (FFF) and big area additive manufacturing (BAAM), of composites and thermoplastics is an area of interest in AM.

An important characteristic to be taken into consideration in both new and existing materials is the behavior they exhibit during extrusion. This behavior is informed in part by the rheological properties of the material, especially the viscosity . Rheological properties of various materials used in extrusion-based AM, ranging from commodity plastics such as acrylonitrile butadiene styrene (ABS) and polylactic acid (PLA) to high-temperature plastics such as polyphenylsulfone (PPSU) and poly (ether ketone ketone) (PEKK) have been studied [2- 6]. The polymer melts used in FFF and BAAM display shear thinning behavior, where the viscosity decreases as shear rate increases [7]. Rheological properties of AM feedstock can be measured using either a parallel-plate rheometer or a capillary rheometer[8, 9].

Rheological properties of AM feedstock can be measured using either a parallel-plate rheometer or a capillary rheometer[8, 9]. On a parallel- plate rheometer, the sample is put between two plates and subjected to oscillatory shear at a specified frequency or frequency range, as shown in Figure 1’s model, while controlling or measuring three factors: angular velocity, angular displacement or torque. These values are then used to find the rheological properties of the material such as the storage modulus, loss modulus, and complex viscosity [9]. A capillary rheometer on the other hand extrudes the molten plastic material through a die, and the force or pressure needed to extrude the material at a given rate is measured [9].An illustration of the capillary rheometer can be seen in Figure 2. The force or pressure and extrusion rate are then used to determine the rheological properties of the material such as the shear viscosity (reference) [9].

The force or pressure and extrusion rate are then used to determine the rheological properties of the material such as the shear viscosity (reference) [9]. Parallel plate rheometers are best suited for low angular frequencies (0.01-628 rad/sec) and capillary rheometer are best suited for high shear rates (1000-10000 1/s) [9]. These two ranges can be correlated via the Cox-Merz rule for some polymers [10].[1]1102As the need for high-performance materials for AM with high specific modulus rises, so has the interest in carbon fiber reinforced polymers (CFRP) [11]. CFRP come in many differentFigure 1. Model of parallel plate rheometerFigure 2. Model of capillary rheometershapes from manufacturers, such as sheets, filament or pellets. Pelletized feedstock is of interest to the authors as that is used in BAAM. Measuring the rheological properties of CFRP comes with different challenges depending on the type of rheometer used.

Pelletized feedstock is of interest to the authors as that is used in BAAM. Measuring the rheological properties of CFRP comes with different challenges depending on the type of rheometer used. In a parallel plate rheometer, it is suggested to use a gap of ten times larger than that of the longest dimension of the filler [12]. In pelletized feedstocks for BAAM, there can be fibers ranging from 50μm to 1000μm, this would require up to a gap of ten millimeters, which is far too large of a gap for a test to be conducted [4]. Pelletized AM feedstock tend to have some initial fiber alignment in the direction in which they were extruded [13]. When these randomly oriented fiber reinforced pellets, are loaded randomly onto the parallel-plate rheometer for testing, the CFRP will have random fiber orientation before and after they have been melted, as illustrated in Figure 3.

When measuring complex viscosity data, the authors observed a large scatter in the data that is hypothesized to be a result of the random orientation of fibers in the sample during loading. Complex viscosity data of fiber reinforced thermoplastics from the parallel plate rheometer do not align with steady shear viscosity from the capillary rheometer data as it does for neat thermoplastics, which could also be due to differences in fiber alignment [14].Figure 3. (from left to right) model of carbon fiber reinforced pellet, top-down view of a plate loaded with carbon fiber reinforced pellets, top-down view of melted random sample1103In this work, the authors discuss the effects of applying a large shear strain to the sample before measuring the viscosity in parallel plate rheology, referred to here as “pre-shearing”. The goal of this work is to determine if and how pre-shearing influences complex viscosity data.

The goal of this work is to determine if and how pre-shearing influences complex viscosity data. Initial findings are reported using complex viscosity data of carbon fiber reinforced ABS recorded by the parallel plate rheometer on pre-sheared samples and compared with samples that have not been pre-sheared. The authors will also compare viscosity data from pre-sheared samples to capillary rheometer data for the same material.Experimental The material used for the experiments was ABS with 20% carbon fiber by weight inpelletized form. The tests were conducted on a Discovery Hybrid Rheometer 2 (DHR-2) from TA Instruments with 25 mm disposable aluminum plates. The DHR-2 uses the upper plate to apply shear to the sample and the lower plate is held stationary. The first step of the experimental process was to determine the linear viscoelastic region, the range of strains where the samples’ structure is not damaged by the experiment, of the carbon fiber reinforced ABS.

The temperature selected for testing was 250°C, a common temperature for processing for 20%CF ABS, and the frequency chosen was 10 radians/second, an angular frequency similar to extrusion shear rates[15]. This established the strain and temperature that would be used for all the experiments: 0.1% and 250°C.After the linear viscoelastic region was determined, the frequency sweeps were performed. For frequency sweeps, the aforementioned temperature and strain amplitude were used to collect complex viscosity data at a range of angular frequencies. Frequency sweeps were carried out on five samples from 0.1 rad/sec to 628 rad/sec. Each sample was loaded by depositing pelletized feedstock onto the plate with random orientation. The loading and melting of the sample can be seen in Figure 4.Figure 4.

Each sample was loaded by depositing pelletized feedstock onto the plate with random orientation. The loading and melting of the sample can be seen in Figure 4.Figure 4. (from left to right) Pellets loaded in melt ring, pellets after melting, before testing, different view of pellets after melting, before testingThe next task was to determine a pre-shearing process that would be used as acomparison for the samples with no pre-shear. The pre-shearing process is defined as loading the1104pellets, melting them and lowering the upper plate to a chosen gap, and rotating the upper plate at a certain angular velocity for a set time prior to performing a typical frequency sweep on a sample. This process is depicted by the model in Figure 5. This process and variables were chosen after experimenting with numerous variations in displacements, velocities and other variables.

This process is depicted by the model in Figure 5. This process and variables were chosen after experimenting with numerous variations in displacements, velocities and other variables. The process is hypothesized to align the fibers with the direction of the shear they are subjected to, and this alignment of fibers would give data more relevant to extrusion, because fibers are aligned along the flow direction during extrusion [13]. Using the pre-shear process, the complex viscosity data of five samples was collected at 250°C and 0.1% strain. Viscosity data was also collected from a 3D printed spiral made from the same material, which would have high fiber alignment circumferentially, to compare the pre-shear process to a known high alignment sample.Figure 5.

(from left to right) Model of both plates with melted random sample, model depicting the rotation of the upper plate to shear the sample, hypothesized model where fibers are given some alignment from the pre-shearing processData was also collected from the capillary rheometer and corrected for non-Newtonian flow using the Rabinowitsch correction and for excess pressure losses due to end effects using the Bagley correction [16, 17]. The Bagley correction was done by collecting data from 3 different die lengths and extrapolating what the viscosity data would be if the die length were zero, where there would be no pressure loss.Results and DiscussionThe data collected from the non-pre-sheared samples were all plotted together, shown inFigure 6. This graph shows the scatter that is often seen by the authors of this study when measuring complex viscosity of CFRP.

This graph shows the scatter that is often seen by the authors of this study when measuring complex viscosity of CFRP. This is an example of how getting repeatable and reliable data can be a challenge if there is no process to normalize or account for the scatter. Repeatable and reliable data will allow for more accurate comparison of data and possibly more meaningful insights about the data. The complex viscosity data from the five data sets are averaged, as shown in Figure 7, but even the averaged data is only representative of the viscosity data that would be collected from a sample with random fiber orientation. This data is less relevant to the extrusion process when the fibers are highly aligned. Therefore, moving forward, the task is to lower the scatter in the data and acquire data similar to extrusion from parallel plate rheology.1105Figure 7.

Therefore, moving forward, the task is to lower the scatter in the data and acquire data similar to extrusion from parallel plate rheology.1105Figure 7. Average viscosity data from the viscosity data of 5 randomly loaded samples (20% CF ABS, 250°C, Air, 0.1% strain)This may raise curiosity as to why capillary rheometer data is not used instead. This is because there is need for viscosity measurements at a wide range of shear rates and complex viscosity can still be used for determining appropriate processing conditions, such as a rise in temperature to lower the complex viscosity or complex viscosity increase from adding fibers to a polymer.Figure 6. Scatter seen in viscosity data of randomly loaded samples (20% CF ABS, 250°C, Air, 0.1% strain)1106The data from the five pre-sheared samples was also plotted to show the scatter of their recorded viscosity, this is displayed in Figure 8.

The scatter of the normal samples with no pre- shear is also shown by the shaded area on the graph. The scatter in the data of the pre-sheared samples is lower by a minimum of 2.5 times and an average of 12.5 times, calculated via subtracting the lowest complex viscosity values (η*) from the highest complex viscosity values and comparing the values of the pre-sheared and non-pre-sheared samples, as seen in Equation 1. The process has provided more repeatable data for complex viscosity measurements of CFRP in parallel plate rheology. The pre-shearing process does have an influence on the scatter of viscosity data, a primary goal of this work.Figure 8.

The pre-shearing process does have an influence on the scatter of viscosity data, a primary goal of this work.Figure 8. Graph of viscosity data of five presheared samples, overlayed with the range of viscosity of normal samples (20% CF ABS, 250°C, Air, 0.1% strain)𝜂𝜂𝑢𝑢𝑢𝑢𝑢𝑢𝑢𝑢𝑢𝑢,𝑛𝑛𝑛𝑛𝑢𝑢𝑛𝑛𝑛𝑛𝑛𝑛−𝜂𝜂𝑛𝑛𝑛𝑛𝑙𝑙𝑢𝑢𝑢𝑢,𝑛𝑛𝑛𝑛𝑢𝑢𝑛𝑛𝑛𝑛𝑛𝑛𝜂𝜂𝑢𝑢𝑢𝑢𝑢𝑢𝑢𝑢𝑢𝑢,𝑢𝑢𝑢𝑢𝑢𝑢−𝑠𝑠ℎ𝑢𝑢𝑛𝑛𝑢𝑢𝑢𝑢𝑒𝑒−𝜂𝜂𝑛𝑛𝑛𝑛𝑙𝑙𝑢𝑢𝑢𝑢,𝑢𝑢𝑢𝑢𝑢𝑢𝑠𝑠ℎ𝑢𝑢𝑛𝑛𝑢𝑢𝑢𝑢𝑒𝑒 = 𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆𝑆 𝑅𝑅𝑆𝑆𝑆𝑆𝑅𝑅𝑅𝑅It may also be observed from Figure 8 that the range of pre-sheared complex viscositydata lies near the lower limit of the normal complex viscosity range. This led to a comparison of the averaged data from the two sample sets, which can be seen in Figure 9. The complex viscosity data of the pre-sheared samples was on average 30% lower than the non-pre-sheared sample, which is another way the pre-shear process affects the viscosity data from the 20% CF- ABS.

This is thought to be due to shear alignment of the fibers in the sample, as hypothesized in the experimental section of this paper, since aligned fibers would lower viscosity and provide more process relevant data for extrusion [13]. This needs to be confirmed via microscopy or some other type of fiber orientation measuring process. This led to interest in how the pre- sheared data would compare to a sample with highly circumferentially aligned fibers, which is the direction of the shear from the rheometer. To compare the two, a 25mm spiral was printed from the same material, which, due to the extrusion aligning fibers in the flow direction, would have circumferentially aligned fibers [13]. This comparison can also be seen in Figure 9 along with an illustration of the printed spiral.

This comparison can also be seen in Figure 9 along with an illustration of the printed spiral. The complex viscosity data of the printed sample aligned within 1% of the averaged data from the pre-sheared samples, giving more information1107[1]suggesting that shear alignment of fibers may be taking place during pre-shear. This added information raised interest in how the pre-sheared viscosity would compare to capillary rheometer data for the same material, as the normal data had not aligned well with steady shear viscosity data from the capillary rheometer. The steady shear viscosity data from the capillary rheometer would have high fiber alignment as fibers align the direction of flow prior to extrusion [13]. These results are plotted together with the averaged complex viscosity data in Figure 9, and the steady state shear viscosity aligned more with the pre-sheared data than the normal data.

These results are plotted together with the averaged complex viscosity data in Figure 9, and the steady state shear viscosity aligned more with the pre-sheared data than the normal data. The alignment of the pre-sheared complex viscosity results with the steady shear viscosity data from the capillary rheometer will be studied further in the future, but for now it acts as another indicator that the pre-shear process can be used to orient the fibers prior to small amplitude oscillatory measurements on a parallel-plate rheometer. Pre-shearing lowers viscosity and increases fiber alignment and these results correspond to what is observed from capillary rheometer data. The reason why the pre-shear process lowers viscosity must be further investigated in the future.Figure 9.

The reason why the pre-shear process lowers viscosity must be further investigated in the future.Figure 9. Graph comparing averaged viscosity data to viscosity data from a capillary rheometer and a printed specimen (20% CF ABS, 250°C, Air, 0.1% strain) Conclusion and Future WorkThis work discusses some issues that are encountered in parallel plate rheology whenusing CFRP and the authors approach to some of these issues, namely the pre-shearing process. The pre-shear process lowers the scatter in data when compared to normal samples by an average of 12.5 times, lowers viscosity by an average of 30%, and increases alignment with capillary rheometer data. The experimental results also show that the pre-shearing process may align the fibers, which will need to be studied more in the future via measuring fiber alignment in both the1108normal and the pre-sheared samples and comparing the values.

There will also be more research into the alignment of capillary rheometer data and pre-sheared parallel plate rheometer data, as this could allow prediction of behavior at high shear rates without running capillary experiments, which is currently done with the Cox-Merz rule for neat polymers.AcknowledgementsResearch sponsored by the U.S. Department of Energy, Office of Energy Efficiency andRenewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle, LLC.11091.2.3.4.5.6.7.8.9.10.11.12.13.14.15.16.ReferencesAjinjeru, C., V. Kishore, J. Lindahl, Z. Sudbury, A.A. Hassen, B. Post, L. Love, V. Kunc, and C. Duty, The influence of dynamic rheological properties on carbon fiber-reinforced polyetherimide for large-scale extrusion-based additive manufacturing. The International Journal of Advanced Manufacturing Technology, 2018. Bertin, M.-P., G. Marin, and J.-P.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEffect of Shield Gas on Surface Finish of Laser Powder Bed Produced PartsColt Montgomery*, Christopher Farnin**, Greg Mellos*, Michael Brand*, Robin Pacheco*, and John Carpenter*Los Alamos National Laboratory, Los Alamos, NM 87545**Department of Materials Science and Engineering, Lehigh University, Bethlehem, PA 18015AbstractAdditive manufacturing (AM) of metals is a novel manufacturing technique that allows for net-shape or near net-shape parts to be produced quickly. Within additive manufacturing a large concern is the produced surface finish, especially for upward and downward facing surfaces on complex geometries. Surface finish is of utmost importance for many engineering applications. In melting of powders, the gas used dominates the thermal conductivity of the metal powder.

Surface finish is of utmost importance for many engineering applications. In melting of powders, the gas used dominates the thermal conductivity of the metal powder. Manipulation of the type of shield gas may provide a means to modify the surface finish without adjustment of established lasing parameters and thereby produce a higher quality part with minimal post processing. These results have potential applications in aerospace, automotive, and biomedical sectors where surface finish requirements coupled with complex geometries are extremely common.IntroductionSurface finish is an important consideration of any part that can be subjected to fatigueloading as it can dramatically affect part life. With traditional manufacturing surface finish control is critical to reduce friction and noise. Additively Manufactured (AM) produced parts are no different with the ability to control surface finish being critical.

Additively Manufactured (AM) produced parts are no different with the ability to control surface finish being critical. AM parts, particularly those of Laser Powder Bed Fusion (L-PBF) are commonly rougher than machined counterparts and this is inherent to the process.The downward facing surfaces are often rough in L-PBF produced parts due to overmelting and a limited heat conduction path provided by the metal powder that the surface is being deposited upon [1] [2]. It has been found that the thermal conductivity of metallic powder is dominated by the thermal conductivity of the gas that fills the pores between the powder particles [3] [4]. This has particular implication to powder-bed based additive manufacturing where overhanging and edge features are built on top of or bordering powder.Strano et al investigated the effect of powder particles on the surface finish of L-PBFproduced parts [5].

Little research has been done, however, in how the inert gas selection affects the resulting surface finish. Shield gas research is traditionally constrained to traditional welding technologies. It has long been known that the shield gas can affect melt pool morphology as well as stability in the traditional welding process [7] [8] [9]. The effect on morphology should be further enhanced within the powder bed process when the results are coupled with the effect of the metal powder. The hypothesis is that by adjusting the shield gas used surface finish can be improved in the laser powder bed fusion process.438MethodsThe artifact chosen represents a variety of surface geometric features found within actualparts. The samples resemble those produced in literature but with added downward facing features such as the arch and horizontal flat [10]. An illustration of the part can be seen in Figure 1.

The samples resemble those produced in literature but with added downward facing features such as the arch and horizontal flat [10]. An illustration of the part can be seen in Figure 1. The sample has angles ranging from 45 degrees up to 90 degrees in 15 degree increments that represent both upward and downward facing surfaces. The angles were chosen to represent the limits of unsupported surfaces in the laser powder bed fusion process. Each side is 5 mm or longer and is 3 mm wide to allow for multiple passes by the chisel probe tip specified by the ASTM B946-11 standard for surface finish on powder metallurgy produced products. This standard corresponds to the same standard used by the machine manufacturers and other researchers in the field [6] [11] . A DektakXT (Bruker, Billerica, Massachusetts) was used to take these measurements using a 12.5 µm ball tip with 3 measurements of 2.5 mm each performed on each surface.

A DektakXT (Bruker, Billerica, Massachusetts) was used to take these measurements using a 12.5 µm ball tip with 3 measurements of 2.5 mm each performed on each surface. It should be noted that since this is contact profilometry it is possible that the tip filtered the data by not being able to fit into extremely narrow valleys. This could be rectified by using a non-contact method in the future. No additional filtering was done on the data.Figure 1: Surface finish sampleThe gases chosen to be tested are argon (standard), nitrogen (standard), and a 75% Argon25% CO2 weld gas mix (non-standard). These three gases are chosen due to their different thermal conductivities and use in traditional welding. Table 1 shows the degree of variation of thermal conductivity for each of the proposed gases as well as the effective thermal conductivity of the powder based on the Yagi-Kunni model for powders [4].

Table 1 shows the degree of variation of thermal conductivity for each of the proposed gases as well as the effective thermal conductivity of the powder based on the Yagi-Kunni model for powders [4]. Nitrogen has the greatest thermal conductivity of the three gases tested by a factor of 1.4. An asterisk designates that the conductivity was calculated as a volumetric ratio of the individual gases thermal conductivities.439Table 1: Shield gas thermal conductivities and effective thermal conductivities ofmetallic powder.Shield GasThermal Conductivity@ 300 K (W/mK)Effective Thermal Conductivity of Powder Bed @ 300 K (W/mK)Argon0.0180.1752775% Argon 25% CO20.0177*0.17241*90% Argon 10% Helium0.03187*0.30524*Helium0.1511.26766Nitrogen0.0260.25076The use of helium gas was attempted but due to the low molecular weight and highintrinsic speeds, the gas was difficult to pump and establish a differential pressure within the equipment.

No build was successfully completed with the helium gas, a mixture is currently being tested of 90% Argon and 10% Helium to take advantage of the high thermal conductivity of helium but maintain nearly the density of argon. In traditional welding a 75% Argon and 25% CO2 gas mixture is commonly used to improve surface finish of stainless steel welds resulting in a wider but shallower melt pool. This tests compositional differences in gas since the mixture has an almost identical thermal conductivity to that of pure argon.The material chosen for these experiments was 304 Stainless Steel (composition is listedin Table 2) because of its widespread use within industry as well as its stability. 304 Stainless steel has applications in the medical, culinary, and general machinery industries. Stainless steel should not have any issues with measurement probe stylus (usually made of diamond) potentially scratching the surface.

Stainless steel should not have any issues with measurement probe stylus (usually made of diamond) potentially scratching the surface. This can be an issue with softer materials because of the force applied to the tip during measurement [12]. Powder for these experiments is 304L Stainless Steel produced via gas atomization and provided by LPW falling within a size range of 15-45 µm and composition listed in Table 2.Table 2: Composition of 304 Stainless Steel PowderFeCrNiMnSiCSPNO69.818.419.561.520.580.0110.0020.010.0550.03815 samples scattered throughout the build area were measured to alleviate locationdependence on the surface finish results. 36 total samples were produced in order to allow for future testing and to further define any location dependence. These samples are distributed440evenly about the build plate so as to allow for any discrepancies due to build plate location or orientation as a result of lasing or gas flow.

These samples are distributed440evenly about the build plate so as to allow for any discrepancies due to build plate location or orientation as a result of lasing or gas flow. The gas flow will change with the density of the gas used, however, for this study the gas flow is not modified from test to test. Location and orientation dependence work is not discussed in this paper but is ongoing.ResultsThe measured values along with box and whisker plots of the data for each gas can beseen in Figure 2 with the argon appearing in green, nitrogen in blue, and the 75/25 mix in yellow. Each dataset consists of 45 measurements of the surface in question.

Each dataset consists of 45 measurements of the surface in question. The Average roughness values are reported to the nearest micron in Table 3 for all three angles and gases.Figure 2: Box and whisker plots along with data for the surface roughness of varioussurface angles and various shield gases441Table 3: Reported average surface roughness measurements as a function of gas andangle of surface. Surface Roughness (µm)Argon Nitrogen 75% Ar 25% CO245 23 23 230 10 9 1145 20 19 24The 75% Argon 25% CO2 mix had issues with oxygen due to the plasma dissociating the carbon and oxygen. This resulted in an increase in measured oxygen levels in the build chamber. Ongoing work is being done to determine if the resulting carbon affected the material composition.DiscussionThe shield gas conductivity seems to have a negligible effect for gases with only a small difference in conductivity.

A larger thermal conductivity difference may produce different results.Nitrogen and Argon have similar Ra values, even though the thermal conductivities aredifferent by a factor of 1.4. This would indicate that a relatively small change in thermal conductivity (on the same order of magnitude) does not greatly change the surface finish. If the downward facing surfaces had been significantly better for nitrogen, as hypothesized, the conductivity would likely be the cause, since all other parameters were held constant.The Argon/CO2 mix had the worst surface finish of the gases tested. This is in contrast towhat was predicted of a better upward facing surface finish because of the carbon addition. The could be due to carbon absorption during deposition leading to an increase in martensite phase fraction and a rougher surface morphology. This result would indicate that gas composition is of a greater importance than the actual thermal conductivity of the gas itself.

This result would indicate that gas composition is of a greater importance than the actual thermal conductivity of the gas itself. To test this the sample will be cross-sectioned and examined for compositional and microstructural changes. The distribution is similar in range for the first and third quartiles for all three gases tested, albeit with different values. This shows that the surface finish distributions for a given angle are independent of shield gas but that the mean of the distribution is gas dependent.ConclusionsBased on current testing there appears to be little difference in the mean surface roughness of the downward facing surfaces with changing of the inert gas used. This agrees with the current hypothesis in literature that downward surface finish is dominated mostly by powder morphology. There could still be a location dependence that is not being seen in the average results.

There could still be a location dependence that is not being seen in the average results. Ongoing work is being done to determine any such dependence.Upward facing surface finish is affected by the gas composition more than the gas conductivity. This intuitively makes sense since the bulk metal conductivity is orders of442magnitude greater than that of the gas conductivity and for this part orientation the powder will be unable to insulate the weld to as great a degree. This result is likely material dependent and would be different for an alloy more sensitive to composition.Future WorkFuture work is to investigate location and orientation dependence of the samples withinall of the builds. This will better inform users where to place parts for optimal surface finish given the gas flow regimes throughout the equipment. The remaining surfaces will also be measured and compared to the current results.

The remaining surfaces will also be measured and compared to the current results. Higher conductivity gases will be tested to determine conclusively their effect on surface finish. Finally the samples will be sectioned and mounted to determine any changes in the microstructure due to the shielding gas for identical process parameters. If successful a valve system could be implemented into the gas line to allow for purging of different shield gases in-situ to selectively change the part material properties throughout the build without modifying lasing parameters.Special ThanksSpecial thanks to Kevin Baldwin of the Center for Integrated Nano-Technologies at LosAlamos National Laboratory for the use of lab equipment for profilometry.References[1] W. E. King, A. T. Anderson, R. M. Ferencz, N. E. Hodge, C. Kamath, S. A. Khairallah and A. M.

fect of Si on the SLM processability of IN738LCR. Engeli*†, T. Etter*, F. Geiger*, A. Stankowski*, K. Wegener†ALSTOM (Switzerland) Ltd, Brown-Boveri Str. 7, 5400 Baden, Switzerland † Institute of machine tools and manufacturing (IWF), ETH Zürich, SwitzerlandAbstractSelective laser melting of high gamma-prime strengthened superalloys such as IN738LC is of interest in stationary gas turbine applications. Differences have been obtained for the hot cracking susceptibility of different powder batches during SLM processing and indications were found that also minor elements influence the SLM processability. By processing a specific powder batch blended by different amounts of pure silicon, the detrimental effect of this element during SLM could be shown.

By processing a specific powder batch blended by different amounts of pure silicon, the detrimental effect of this element during SLM could be shown. Therefore, the control of this minor element is crucial to decrease the hot cracking tendency and can extend the SLM processing window of this alloy.IntroductionSelective Laser Melting (SLM) is an emerging additive manufacturing method for the fabrication of complex parts out of metal powder. The process consists of repeated application of thin (20-60 μm) powder layers which are subsequently densified by a scanning laser beam according the layer information from a sliced 3D model. This process provides new possibilities in the design of complex components and is therefore of large interest for applications in stationary and aeronautical gas turbines. For such applications, high temperature alloys such as nickel based superalloys are used for components in the hot gas path of gas turbines.

For such applications, high temperature alloys such as nickel based superalloys are used for components in the hot gas path of gas turbines. The high strength of this alloy class is based on a strong solid solution strengthening of the fcc γ-Ni matrix and on precipitation hardening by ordered, coherent intermetallic phases such as γ´-Ni3(Al,Ti) or γ´´-Ni3Nb. Further strengthening is provided by solid solution strengthening of the intermetallic phases and grain boundary strengthening with carbides, boron and zirconium. This alloy class is known to be difficult to process by SLM and many alloys are susceptible to severe hot cracking as reported by Gasser et al. [1] and Kelbassa et al. [2] for the alloy MarM247, by Uhlmann and Urban [3] for René80 and by Carter et al. [4,5] for CM247LC. As shown by Tomus et al. [6], this hot cracking problematic is not limited to γ´-strengthened alloy, but also exists for the solid solution strengthened alloys such as Hastelloy X.

[4,5] for CM247LC. As shown by Tomus et al. [6], this hot cracking problematic is not limited to γ´-strengthened alloy, but also exists for the solid solution strengthened alloys such as Hastelloy X. In their work, Tomus et al. showed that the minor elements Mn and Si affect the hot cracking susceptibility. In a recent work of the authors [7], the SLM processability of various batches of IN738LC was assessed and it was found that also for this γ’-strengthened alloy, the different batches strongly differ in their hot cracking susceptibility. Though being within the original specification for IN738LC given in Table 2, strong variations in the crack densities up to factor of ten were found in the different batches (see Figure 1).

Though being within the original specification for IN738LC given in Table 2, strong variations in the crack densities up to factor of ten were found in the different batches (see Figure 1). Indications were found that the minor element Si has a strong effect on cracking susceptibility even on a very low content level, but an unambiguous identification of the critical elements was not possible due to the large amount of varying elements compared to the number of powder batches which were investigated.It is already known from conventional welding research that some elements such as boron, sulfur orphosphorous strongly affect hot cracking even on ppm level [8]. To unambiguously identify the influence of the different elements, either a statistical evaluation of a large number of different batches (>> number of varying elements) is required or test series must be performed where only the content of specified elements varies while the others are kept constant.

The latter might be achieved by atomization of different batches from a common master melt with different additions of a specific element. However, due to the high atomization cost for such alloys both of these ways are very cost intensive.Figure 1: Representative microsection of a SLM made sample processed under equal conditions from a) a ‘good’ and b) a ‘bad’ IN738LC powder batch.Another approach to assess the influence of minor elements is to blend a powder batch with low amounts of a second powder enriched with the element to be investigated. Such an approach is more versatile and cost effective, as it allows the preparation of various compositions from the same source powders. However, it requires that both components are completely melted and well homogenized during the SLM process. In a first step, the homogenization of a simple powder blend consisting of 50% IN738LC and 50% CM247LC powder during SLM was assessed.

In a first step, the homogenization of a simple powder blend consisting of 50% IN738LC and 50% CM247LC powder during SLM was assessed. These alloys were chosen because they can be well distinguished by backscatter electron microscopy and by EDX mapping, allowing an easy investigation of the homogenization efficiency during SLM. In a second step, this blending approach was applied to investigate the effect of Si on the processability of IN738LC.Experimental procedureIn the predecessor experiment, the capability of the SLM process to homogenize a powder blend was assessed by the investigation of samples made of a 50wt%:50wt% powder blend of IN738LC and CM247LC. Both powders were gas atomized and mainly consist of spherical particles with a D50 of 18 μm (IN738LC) and 33 μm (CM247LC) respectively. The chemical composition of the source powders is given in Table 1.

The chemical composition of the source powders is given in Table 1. To prepare the powder mixture, 500g of each powder was added to a plastic bottle and blended in a tumbling mixer for 20 minutes.To investigate the effect of Si on the SLM processability of IN738LC, four different blends of IN738LCpowder with A) 0 wt-%, B) 0.05 wt-%, C) 0.125 wt-% and D) 0.2 wt-% pure Si powder were prepared. The IN738LC powder showed a spherical morphology with a D50 of 36 μm. An irregular shaped, fine Si powder with a D50 of 3.6 μm was used for blending. The chemical composition of the two source powders is shown in Table 2. After adding the appropriate amount of Si powder to the IN738LC powder, the blends were homogenized in a tumbling mixer for 10 minutes and further de-agglomerated by ball milling for 120 min.

After adding the appropriate amount of Si powder to the IN738LC powder, the blends were homogenized in a tumbling mixer for 10 minutes and further de-agglomerated by ball milling for 120 min. To ensure equal conditions, the same procedure was also applied to the pure IN738LC powder batch.In order to cope with the low amount of powder available, a low volume insert developed by Inspire AG was used. From each powder blend, three cuboids with dimensions 15x10x15 mm were prepared under equal conditions using a parameter set developed by Rickenbacher et. al. [9] for IN738LC using a layer thickness of 30 μm. A crosswise alternating island based scan strategy was applied with scan direction ±45° to the edges of the samples. To investigate the crack surface, a special specimen geometry was fabricated which is shown in Figure 2. The specimen was designed such that it can be easily cleaved by using a bench vise.

To investigate the crack surface, a special specimen geometry was fabricated which is shown in Figure 2. The specimen was designed such that it can be easily cleaved by using a bench vise. For this specimen, the same process parameters are used, but a bidirectional scan strategy was applied as explained in Figure 2. This set-up allows the investigation of the surface of both cracks running parallel and perpendicular to the weld lines.Figure 2: Schematic of the test geometry used to investigate the crack surfaces. A bi-directional scan strategy is used where the scan vectors on one side of the specimen run parallel and on the other side perpendicular to the later fracture plane.

A bi-directional scan strategy is used where the scan vectors on one side of the specimen run parallel and on the other side perpendicular to the later fracture plane. After forced fracture of the samples, the crack surface of cracks running parallel and perpendicular to the scan line can be investigated on the corresponding side of the specimen.For metallographic analysis, two xy-microsections perpendicular to the build-up direction were prepared from each sample by grinding with a 150 grit disk (corundum) followed by polishing with 15 µm, 9 µm, 6 µm, 3 µm, and 1 µm diamond suspension. The porosity and the crack density were evaluated by quantitative image analysis using a plugin created by the first author for the software ImageJ (http://imagej.nih.gov/ij/, [10]). For each sample, 8 images with 50x magnification were evaluated.

For each sample, 8 images with 50x magnification were evaluated. The porosity, the crack density and the corresponding uncertainties (95%) reported in this work were calculated from the average values determined for the three samples of each powder batch. A Tescan MIRA3 XMU FEG-SEM equipped with an AZtecEnergy EDX system from Oxford Instruments, operated at 20kV and with a nominal beam current of roughly 1 nA, was used for the investigation of polished microsections and EDX mapping. All other SEM images were made with a JEOL JSM-6010LA scanning electron microscope. Thermodynamic properties such as solubility limits and melting temperature were calculated with the software ThermoCalc V4 using the database NIDATA V4.Results & DiscussionTo assess the capability of the SLM process to homogenize powder blends, specimen prepared out of a50wt-% IN738LC:50wt% CM247LC powder blend were investigated by backscatter electron (BSE) microscopy and EDX mapping.

These two alloys differ strongly in several elements and can therefore be well distinguished by their z-contrast in the BSE images or by EDX mapping the Cr-Kα signal. In Figure 3, such a BSE image and an EDX map from a polished microsection of a SLM processed sample is shown.Table 1: Chemical composition of the predecessor powders and of theoretic and measured (EDX) composition of samples processed out of the IN738LC/CM247LC powder blend.NiCrCoMoW TaAlTiHfCBZrNbFeO2N2IN738LCSupplier bal.16.28.61.82.71.73.53.30.140.0110.0230.920.82400*9*CM247LCSupplier bal.8.09.10.59.93.25.60.71.30.060.0130.0090.05145*117*calculatedbal.12.18.81.16.32.54.62.00.70.100.0120.0160.460.4327363EDX58.912.38.91.37.12.94.72.10.50.500.42* LECO analysis[wt-ppm]Predecessor powder50/50 mixturePowder blend[wt-%]The slight contrast visible in the BSE images originates from channeling effects due to different grain orientations.

However, no evidence for incomplete solution of the different powder particles was found in the investigated microsection. This is supported by the EDX map of the Cr-Kα signal which shows a very uniform distribution of the element. For the given process conditions, the mixing effect within the 100-200 μm broad melt pool is strong enough to homogenize the powder blend despite the rather large particles ranging between 8to 35 μm for IN738LC and 11 to 57 μm for CM247LC. The large mixing effect might be caused by strong Marangoni convection within the melt pool which can be in the range of several meters per seconds [11,12].

The large mixing effect might be caused by strong Marangoni convection within the melt pool which can be in the range of several meters per seconds [11,12]. As shown in Table 1, the resulting chemical composition obtained by EDX analysis matches well the expected composition calculated from the composition of the source materials.Figure 3: a) Backscatter electron microscopy image and b) EDX map for Cr-Kα of a polished microsection of a SLM processed sample from a blended IN738LC/CM247LC powder mixture.The findings from this preliminary experiment encourage the assumption that the effects of alloying elements on SLM processing can be investigated by processing powder blends. This approach will have its limitations, e.g. when blending powders with very different melting points or using different process conditions resulting in other melt pool dynamics.

This approach will have its limitations, e.g. when blending powders with very different melting points or using different process conditions resulting in other melt pool dynamics. Also the solubility of the elements in the matrix can be critical, even though the rapid solidification occurring in the SLM process might extend the solubility limits. These limits must be investigated in a future work.Based on this promising result, the powder blending approach was applied to verify the detrimentaleffect of Si found in the author’s prior work [7]. Four blends of IN738LC with 0, 0.05, 0.125 and 0.2 wt-% Si powder were prepared and investigated. The expected chemical composition of the blend can be found in Table 2. The element of interest for this investigation, silicon, has a melting point of 1414°C, only around 75°C higher than the liquidus of IN738LC.

The element of interest for this investigation, silicon, has a melting point of 1414°C, only around 75°C higher than the liquidus of IN738LC. The maximum solubility of Si in pure Ni is around 1.5 wt-%, considerably higher than the range to be investigated which is below 0.2 wt-%. Based on these two considerations, it is expected that a good homogenization should also be obtained for powder blends with Si.Table 2: Chemical composition of the used powders used to investigate the effect of Silicon on SLM processability. In the first two rows, the nominal composition of IN738LC according the technical datasheet “Alloy IN-738LC: Technical data” [13] is shown. The next two rows show the chemical composition of the source powders according to the supplier’s datasheets. The four last rows show the expected composition of the blended powder mixtures A-D as calculated from the source powder compositions.

As in the other blends, the silicon particles (marked with arrows) are well distributed and adhere to the surface of the larger IN738LC particles.Because of the low Si contents to be investigated, a very fine Si powder with a D50 of 3.6 μm was used to get a homogenous powder mixture. Assuming that the particles of the two powders all had a uniform size corresponding to their D50 value, a rough estimate of the number fraction of Si particles in the different blends can be calculated using the bulk densities of IN738LC (8.1 g/ccm) and Si (2.33 g/ccm). This rough estimation results in number fractions of Si particles in the blends between 63% for the 0.05 wt-% blend and 87% for the 0.2 wt-% blend. For all blends, more Si particles than Ni particles are present which should favor a good homogenization during SLM. In Figure 4, a BSE image of the 0.05-wt% Si powder blend after de- agglomeration by ball-milling is shown.

In Figure 4, a BSE image of the 0.05-wt% Si powder blend after de- agglomeration by ball-milling is shown. It can be observed that the Si particles are well distributed within the powder blend.Figure 5: Representative microsections of SLM processed samples made of IN738LC blended with a) 0%, b) 0.05%, c) 0.125%, d) 0.2% Si powder.The evaluation of the polished microsections showed a strong increase of the crack density with increasing Si powder fraction as shown in Figure 5. The results from the quantitative image analysis are given in Figure 6. While the porosity is at low values around 0.1% for all blends, the crack density is increasing from around 0.1 mm/mm2 for the pure IN738LC batch up to 1.57 mm/mm2 for the blend with 0.2 wt-% Si powder.Figure 6: Results from quantitative image analysis of the xy-microsections.

While the fraction of Si powder does not influence the porosity, a strong increase of the crack density can be observed with increasing Si powder fraction.The microsections of the SLM processed samples made of pure IN738LC powder and the 0.2 wt-% Si powder blend was investigated by scanning electron microscopy. Figure 7 shows two BSE images of these microsections. The microstructure consists of small grains consisting of cells of around 1μm size and very fine carbides (bright spots in Figure 7) with a size considerably below 250nm. In both samples, some very fine pores are present (dark spots in Figure 7). However, no evidence for undissolved Si particles or Si enrichment could be observed.Figure 7: Backscatter electron images of polished microsections from SLM processed samples out of a) pure IN738LC powder and b) IN738LC powder blended with 0.2 wt-% Si powder.

Fine pores (dark) and carbides (bright) can be seen in all samples, but no sign of remaining silicon particles or enrichment were observed. The contrast of the different areas in the images originates from channeling effects caused by slightly different orientations of the grains.As Si is the only element whose concentration is changed significantly in the different blends (see Table 2), the present findings clearly show the detrimental effect of very small amount < 0.2 wt-% of Si in IN738LC on its SLM processability. In author’s prior work [7], the crack densities of SLM processed specimens made from different gas atomized IN738LC powders varied between 0.3 and 3 mm/mm2. Though the Si concentrations in the present work varied in a similar range, smaller variations in the crack densities were observed.

Though the Si concentrations in the present work varied in a similar range, smaller variations in the crack densities were observed. However, it needs to be noted that these values are not directly comparable as the specimens of the present investigation were processed with different SLM equipment using a process parameter with considerably slower scan speed. For directionally solidified IN738LC, Zhu et al. [14] reported that the segregation of Si increases with increasing solidification rate from 12 mm/h to 50mm/h. If this trend also holds true for the much higher solidification rates present during SLM, it might contribute to the higher crack densities observed in theauthor’s prior work.

If this trend also holds true for the much higher solidification rates present during SLM, it might contribute to the higher crack densities observed in theauthor’s prior work. Furthermore, those different batches varied in many elements simultaneously and the higher crack densities might also be caused by a synergistic effect of different detrimental elements.Figure 8: Secondary electron image of top surface from an SLM processed sample out of IN738LC with 0.2wt-% Si powder addition. As in the other samples, the cracks run both perpendicular and parallel to the scan line (see arrows).All samples reveal cracks open to the top surface, which can run both along the scan line as well asperpendicular to the scan lines (see Figure 8). The SE images in Figure 9 show details of the fracture surface of the specimen made out of pure IN738LC powder. Large areas of the fracture surface exhibit dimple-like features as shown in image Figure 9a, characteristic for a ductile forced fracture.

Large areas of the fracture surface exhibit dimple-like features as shown in image Figure 9a, characteristic for a ductile forced fracture. Figure 9b and Figure 9c show detail images of a crack running along and perpendicular to the scan vectors respectively. Both show rounded, cellular or dendrite like features characteristic for hot cracks forming under presence of a liquid phase [15].Figure 9: Secondary electron images of the crack surface from a broken open specimen processed out of IN738LC with no Si powder addition. a) Fracture surface showing dimples characteristic for a ductile forced fracture. b) Fracture surface of a hot crack perpendicular to the scan direction. c) Fracture surface of a hot crack running parallel in the center of the weld lines.

b) Fracture surface of a hot crack perpendicular to the scan direction. c) Fracture surface of a hot crack running parallel in the center of the weld lines. The smooth, dendrite like features shown in b) and c) are typical for fractures occurring under presence of a liquid phase.Figure 10: Secondary electron images of the crack surface from a broken open specimen processed out of IN738LC with 0.2wt-% Si powder addition. a) Fracture surface of a hot crack running perpendicular to the scan direction. b) View on a fracture surface running parallel to the scan lines. c) Close up of a hot crack running parallel to the scan lines. The smooth, dendrite like features shown are typical for fractures occurring under presence of a liquid phase.The fracture surface of the specimen made out of the 0.2 wt-% powder blend shown in Figure 10 exhibits very similar features, but the area showing the hot crack features (Figure 10 a and c) almost covers the whole fracture surface.

This can be well seen in the overview image (Figure 10b) looking on the fracture surface running parallel to the scan lines. The numerous cleavages in this fracture surface originate from crack running perpendicular to the scan vectors. The present investigation showed that cracks in as-built, SLM processed IN738LC are either solidification cracks, forming at the final stage of solidification, or liquation cracks, forming through liquation of segregated low-melting phases or constitutional liquation of precipitates during reheating by an adjacent scan line. According Wolf [15], surfaces of solidification cracks typically show a pronounced columnar structure, whereas the structure on the surface of liquation cracks is more irregular. It is therefore expected that cracks in SLM processed IN738LC are more likely solidification cracks. This is supported by the fact that the cracks in the top layer are mostly open to the top surface and often run along the weld center line.

This is supported by the fact that the cracks in the top layer are mostly open to the top surface and often run along the weld center line. Furthermore, only very fine precipitates are present in the as-built microstructure due to the very high cooling rates characteristic to the SLM process. As fine particles can dissolve faster upon heating, they are less likely to constitutional liquate as shown by Tancret [16] for liquation of γ` precipitations. Solidification cracking might be aggravated by segregation of Si to the interdendritic zone where it could result in an increased mushy zone as reported by Zhu et al. for casted IN738LC [14]. It is furthermore possible that Si affects the solid-liquid interface energy which could, according the work of Rappaz et al. [17,18], result in repulsive grain boundaries which also promote solidification cracking.

[17,18], result in repulsive grain boundaries which also promote solidification cracking. Further investigations are ongoing to clarify the cracking mechanism observed with increasing Si content.Summary & Outlook Even in low concentrations below 0.2 wt-%, Si increases the alloy’s tendency for hot cracking. Thecrack surfaces revealed that the cracks are formed under presence of a liquid phase, but the exact mechanism behind is not yet clear and requires further investigations. The findings of the present work emphasize that alloy specifications optimized for SLM are necessary, especially if the process is to be industrialized. The creation of such a specification requires understanding of the influence of the different alloys on SLM processing.

The creation of such a specification requires understanding of the influence of the different alloys on SLM processing. Based on the new findings, the IN738LC composition was optimized for the use for SLM (patent pending [19]) and is applied with good success to create parts with very low crack densities.AcknowledgmentThe authors are indebted to the Institute for rapid product development, Inspire AG, especially Mr. Adriaan Spierings for providing access to the SLM equipment and the valuable cooperation.References[1][2][3][4][5]A. Gasser, G. Backes, I. Kelbassa, A. Weisheit, and K. Wissenbach, “Laser Additive Manufacturing,” Laser Technik Journal, vol. 7, no. 2, pp. 58–63, Feb. 2010. I. Kelbassa, P. Albus, J. Dietrich, and J. Wilkes, “Manufacture and repair of aero engine components using laser technology,” in Proceedings of the 3rd Pacific International Conference on Application of Lasers and Optics, 2008, pp. 208–212. E. Uhlmann and K.

FECT OF SINTERING PARAMETERS AND FLOW AGENT ON THE MECHANICAL PROPERTIES OF HIGH SPEED SINTERED ELASTOMERFarhana Norazman and Neil HopkinsonDepartment of Mechanical Engineering, The University of Sheffield, Sheffield S1 3JD United KingdomAbstractHigh Speed Sintering (HSS) is an Additive Manufacturing process that creates parts by sintering using inkjet and infra-red lamp technology rather than laser systems employed in Laser Sintering (LS). This research investigated the effects of machine parameters (sintering power, bed temperature) and the addition of fumed silica flow agent on the tensile properties of thermoplastic elastomer parts processed using HSS. The results showed improved elongation at break values by a factor of more than 2X compared to reported values for LS of the same thermoplastic elastomers.

The results showed improved elongation at break values by a factor of more than 2X compared to reported values for LS of the same thermoplastic elastomers. At constant parameters, improved tensile strength and tensile modulus were observed with the addition of flow agent into the sintering mixture.KeywordsHigh speed sintering, thermoplastic elastomer, tensile properties, sintering power, flow agentIntroductionAdditive Manufacturing (AM) is defined as “the process of joining materials to make objects from 3D model data, usually layer upon layer, as opposed to subtractive manufacturing technologies” (1). There are currently over twenty recognised Rapid Prototyping (RP) processes based on the “additive” principle, though method of layers consolidation may differs from one another.

There are currently over twenty recognised Rapid Prototyping (RP) processes based on the “additive” principle, though method of layers consolidation may differs from one another. Classification is made according to the physical state of raw material used in the process; liquid-based systems (eg: Stereolithography), solid- based systems (eg: Fused Deposition Modelling) and powder-based systems (eg: Laser Sintering, High Speed Sintering) (2). Depending on the type of process and hardware, a range of metal, polymer and ceramic materials can be used with AM to create end-use parts with minimal post-processing.The main advantage of AM is the design freedom that comes with its tool-less and mould-free processes, enabling consumers to produce parts of almost any geometric shape or feature (3). Design customisation is also made easy and cost-effective with AM as it only takes place within the CAD design.

Design customisation is also made easy and cost-effective with AM as it only takes place within the CAD design. This eliminates many restrictions of conventional design for manufacture and assembly (DFMA) such as the need for undercuts, draft angles, reduced part counts, etc and therefore improves the design process as a whole (2).Laser Sintering (LS) is a popular additive process used for making polymer parts. It uses a high power CO2 laser to sinter powder raw material on a surface bed to create desired shape. Nylon-based materials have been widely used with LS for many years in both RP and Rapid Manufacturing, particularly the nylon polyamide (Nylon-12) for its superior properties (4).

Nylon-based materials have been widely used with LS for many years in both RP and Rapid Manufacturing, particularly the nylon polyamide (Nylon-12) for its superior properties (4). Material suppliers such as 3D Systems and ALM specialise in materials development and has manufactured a range of polyamides (eg: filled and unfilled nylon) and elastomers (eg: Thermoplastic elastomer, TPE) to be used with LS (5).High Speed Sintering (HSS) is a relatively new AM process invented at and patented by Loughborough University in the UK, with its process developments detailed by Hopkinson et al (6). It utilises inkjet print head and infra-red technology to manufacture products layer by layer from polymer powder material. The process incorporates the same technique as LS with the exclusion of laser disintegrated into the following; IR lamp acts as the heating source and inkjet print head defines the part geometry.

The process incorporates the same technique as LS with the exclusion of laser disintegrated into the following; IR lamp acts as the heating source and inkjet print head defines the part geometry. HSS has several advantages over LS – reduced machine cost by laser elimination and reduced build time as it sinters layer by layer. In common with all AM processes, HSS is capable of producing complex geometry parts with superior mechanical properties.Application of AM is widespread across aerospace, automotive and consumer industry. LS is used to manufacture air ducts for Boeing F-18 military jets.

LS is used to manufacture air ducts for Boeing F-18 military jets. New Balance, a major footwear manufacturer has adopted LS in producing high-performance customised running shoes for elite athletes.High Speed SinteringFigure 1 shows the High Speed Sintering machine and its key parts; powder bedheater, infrared lamp, roller system, inkjet printhead and powder bed.The build procedure on the HSS machine can be broken down into three main stages; pre-processing, building and post-processing. The pre-processing involved modelling of 3D part on a CAD software (Solidworks) then subsequently sliced into a stack of 2D layer files or slice files (SLC), each of 100μm thickness in preparation for the HSS machine.

The layer files were transferred to the machine where build parameters were set.Once the prepared powder was fed into the feed chamber and build parameters were set, the building process began with deposition of a layer of raw powder onto the part bed by roller action, with an IR preheat stroke to warm up the powder prior to sintering. Monochromatic bitmap images of desired part were printed onto the powder bed using a Radiation Absorbent Material (RAM), which upon heating by IR lamp will absorb thermal energy thus melted and sintered the powder underneath. Unprinted powder will remain unsintered and acted as a support to the build. New layer of raw powder will be redeposit and this process was repeated until all layers have finished building. Blank layers may be included in the build to separate different parts.The parts were left to cool with all heating parts switched off, before the cake was removed from the machine.

Blank layers may be included in the build to separate different parts.The parts were left to cool with all heating parts switched off, before the cake was removed from the machine. Post-processing involved powder removal through bead blasting process to obtain finished parts. Other properties enhancement technique such as infiltration of parts may follow if needed.This research highlights the effects of three main build parameters on the mechanical properties of finished parts, using a TPE material. Previous researches concerning the HSS process include the study of excess powder hardness using Nylon with HSS (6) and the effect of Infra-red power level on sintering behaviour of Nylon-12 (7).Powder bed heaterInfrared lampRoller SystemInkjet printheadPowder bedFigure 1 – High Speed Sintering machine and key partsAccording to Kang (8), sinterability and the sintered microstructure of a powder compact are mainly determined by material variables and process variables.

A set of key HSS build parameters are divided into the two groups and defined in Table 1:Material variables  Layer Thickness – The thickness is powder of determined by the lowering of build platform.eachlayerProcess variables  Build Bed Jacket & Overhead Temperatures – The base and overhead temperatures of bed determines at sintering.the powdertemperaturePowder Ratio – The composition of powder material which may include additives or fillers.Sintering Power – The sintering rate is determined by the sintering speed.Surface Preheat – The rate at which powder is preheated before sintering.Table 1 – Sintering parametersElastomerThermoplastic elastomer (TPE) is elastomeric with the properties of thermoset rubber. It softens when heated to a flowable state but does not cure or set under heat as does thermoset hence its chemical properties were unchanged thus promoting recyclability (9).

It softens when heated to a flowable state but does not cure or set under heat as does thermoset hence its chemical properties were unchanged thus promoting recyclability (9). This study will focus on the manufacturing and testing of TPE210-S, a flexible elastomer sintering material supplied by ALM. The ASTM definition of elastomer is “a material which at room temperature can be stretched repeatedly to at least twice its original length, and, uponrelease of the stress, will return immediately to its approximate original length” (10). The general mechanical properties for TPEs include low modulus and high yield strain.

The general mechanical properties for TPEs include low modulus and high yield strain. Table 2 below listed the main properties of TPE210-S as provided by the manufacturer, ALM when processed using LS.Properties Average Particle Size (D50) Melting Point Tensile Modulus Elongation at Break (uninfiltrated) Shore Hardness (uninfiltrated)Value 85 μm 178 °C 8 MPa 110 % 40Table 2 – TPE210-S Material PropertiesTPE has existed for 40 years and was favourable to conventional rubber for its ease of processibility and flexibility. Injection moulding is the principle process used in the fabrication of TPE parts, followed by extrusion process. Injection moulded TPEs are advantageous of costs at rapid production rates. However in comparison, LS is more economical for low volume part production (less than 14000 unit) (2). TPEs have replaced thermoset rubbers in most of its former applications.

However in comparison, LS is more economical for low volume part production (less than 14000 unit) (2). TPEs have replaced thermoset rubbers in most of its former applications. TPE parts are widely used in automotive applications for both exterior (bumpers) and interior (dashboard) parts, including under-the- hood automotive applications such as front-wheel drive components and air ducts (9).Methodology1. Manufacture of test partsTo investigate the influence of build parameters, test specimens were manufactured with variation in build bed overhead temperature, powder ratio and sintering power. A Design of Experiment was created as followed. Table 3 lists the complete build parameters used throughout the project. Table 4, Table 5 and Table 6 were build parameters with variations in sintering power (speed), build temperature and powder composition (presence of flow agent), respectively.

Tensile testingTensile tests were carried out according to ASTM D638 Standard Test Method for Tensile Properties of Plastics (11) using a Tinius Olsen H5K-S Universal Testing Machine equipped with model 500LC Extensometer and HT36 grip under ambient conditions. The crosshead speeds of 5mm/min and 10mm/min were maintained during the tests. The method provided calculations of the elongation at break (%), young’s modulus (MPa) and ultimate tensile strength (MPa). Figure 3 shows a typical stress-strain curve obtained by the tensile test.Ultimate Tensile StrengthElongation at BreakYoung’s ModulusFigure 3 – EaB, YM and UTS determined from ASTM D638 Tensile Test3. Thermal analysisThe thermal analysis of raw powder material was carried out as outlined by ASTM D3418 Standard Test Method for Transition Temperatures and Enthalpies of Fusion and Crystallization of Polymers by Differential Scanning Calorimetry (11).

The tests were performed using a double furnace PerkinElmer DSC8500 instrument. The powder sample tested were 100% Virgin TPE210-S (0.0113g) and Virgin TPE210-S mixed with 0.2% Cab- O-Sil (0.0167g). The standard reference material was an empty pan.The samples were heated from 20°C to 210°C at the rate of 20°C/min. The temperature was held isothermally at 210°C for 1 min. Then it was cooled back to 20°C at the rate of 20°C/min. The temperature was held isothermally at 20°C for 1 min. The cycle of heating, temperature hold and cooling was repeated at the same rate . The DSC curves were plotted for all samples.4. Surface roughness testingSurface roughness testing was done using a Mitutoyo Surftest contact profilometer rig to quantify the quality of high speed sintered TPE parts surface. Constant speed of 0.5mm/s was maintained and evaluated along 20.0 mm length of the type IV ASTM test specimen during the test. The test was performed on both top and bottom surfaces of specimens.

Constant speed of 0.5mm/s was maintained and evaluated along 20.0 mm length of the type IV ASTM test specimen during the test. The test was performed on both top and bottom surfaces of specimens. Values of roughness parameter; average roughness, Ra and average maximum height, Rz were obtained through this method.Results and Discussion1.

Values of roughness parameter; average roughness, Ra and average maximum height, Rz were obtained through this method.Results and Discussion1. Mechanical Properties a) The effect of sintering speed at two sets of constant temperature and powder composition parameter.Parameter 1Parameter 2)%300( k a e r B250200t a n o i t a g n o l E15010050050607080Sintering speed (mm/s)Figure 4 – Variation of EaB with sintering speedFigure 4, Figure 5 and Figure 6 show the graphs of mechanical properties EaB, UTS and YM plotted against increasing sintering speed at two constant bed temperature 90°C (Parameter 1) and 100°C (Parameter 2), respectively.The maximum value of EaB, UTS and YM are observed at the lowest sintering speed of 50mm/s, which are 265%, 2.20 MPa and 4.91 MPa respectively) a P M( h t g n e r t S e l i s n e T e t a mi t l4.003.002.001.00From it can be figures, concluded that EaB increases with speed.

UTS decreasing increases with decreasing sintering speed, and at a higher rate at low high compared temperature temperature. The value of YM decreases with increasing sintering speed at both set of temperatures. The variation is not linear and remains low across all set of build parameter.thesesinteringtoU0.0050607080Sintering speed (mm/s)Figure 5 – Variation of UTS with sintering speed15.00) a P M12.00(s u l u d o M9.006.00s ' g n u o Y3.000.0050607080Though at different intensities, it can be observed the mechanical properties of HSS printed TPE increase with decreasing speed; highest properties achieved at lowest speed and conversely true. The influence of this parameter on mechanical properties can be related to energy density and sintering behaviour.

According to Andrew Number, established model formulated for LS energy density calculation, a decrease in sintering speed results in an increase in energy density and directly mechanical properties.thatanSintering speed (mm/s)Figure 6 – Variation of YM with sintering speedb) The effect of bed temperature at two sets of constant sintering speed and powder composition parameter.Parameter 1Parameter 2)%( k a e r Bt a n o i t a g n o l E30025020015010050Figure 7, Figure 8 and Figure 9 show the graphs of mechanical properties EaB, UTS and YM plotted against increasing build bed temperature constant sintering speed 70 mm/s (Parameter 1) and 60 mm/s (Parameter 2), respectively.attwo09093100110Build bed temperature (°C)Figure 7 – Variation of EaB with bed temperatureThe maximum value of EaB, UTS and YM are observed at the highest build bed temperature of 110°C, which are 214%, 1.55 MPa and 3.32 MPa respectively.) a P M( h t g n e r t S e l i s n e T e t a m4.003.002.001.00A slight fluctuation can be observed in values of EaB from 90-100°C, which can be regarded as constant before a rapid increase at 110°C.

The same variation exist in tensile properties; an almost horizontal gradient that shows little effect of build bed temperature on UTS and YM until reaching 110°C.i t lU0.009093100110Build bed temperature (°C)Figure 8 – Variation of UTS with bed temperature15.00) a P M12.00(s u l u d o M9.006.00s ' g n u o Y3.000.009093100110largely Build bed determines the powder temperature the bed which are set across in surface. According Figure 13, the temperature range of 100 – 110 °C is where glass transition occurs for TPE210-S. temperature During this critical range, the material softens and flows at low viscosity before it melts completely, which provides the best characteristics for sintering.

It is possible that high degree of sintering with full coalescence took thus place at provided tensile properties of the sintered parts.temperatureto DSCthistemperature,improvedBuild bed temperature (°C)Figure 9 – Variation of YM with bed temperaturec) The effect of flow agent addition at two sets of constant sintering speed and temperature parameter)%( k a e r Bt a n o i t a g n o l E30025020015010050Parameter 1Parameter 2A flow agent (fumed silica) was added to the following two best set of parameters to further enhance their finished parts properties; bed : Parameter temperature, 50 mm/s sintering speed) & Parameter 2 : (110°C bed temperature, 70 mm/s sintering speed).1(100°C00.00.20.51.0Flow agent (% by weight)Figure 10 – Variation of EaB with flow agent levelThe flow agent addition on EaB, UTS and YM are plotted in Figure 10, Figure 11 and Figure respectively, with comparison to 100% virgin-powder sintered parts.influence of124.00) a P M( h t g n e r t S e l i s n e T e t a m3.002.001.00i t lU0.000.00.20.51.0Flow agent (% by weight)Figure 11 – Variation of UTS with flow agent level15.00) a P M12.00(s u l u d o M9.006.00s ' g n u o Y3.000.000.00.20.51.0Flow agent (% by weight)The maximum value of EaB, UTS and YM are observed at 0.2% flow agent by weight, which are 284%, 2.57 MPa 10.90 MPa and respectively.All steady decrease in EaB, UTS and YM when more flow agent is mixed with virgin powder pass 0.2% level.figuresindicate aThe influence of polymer additives on HSS part performance can be explained by Frenkel’s Model of Sintering which relates particle viscosity to predicted sintering rate.

for Zero TPE210-S in unknown. However, it is relatively high compared to Nylon-12. It can be speculated that fumed silica promotes fluidity and lowers viscosity of the TPE210-S mix. A decreased viscosity lowers sintering time, thus allowing less energy addition. As low thermal energy tensile is absorbed and properties depend largely on the this will heat, amount subsequently result in parts of inferior propertiesshear viscosityrateofFigure 12 – Variation of YM with flow agent leveld) Comparison between mechanical properties of parts processed by High Speed Sintering, Selective Laser Sintering and Injection MouldingProcessHSS HSS LS (5) Injection Moulding (9)% by weight of fumed silica mix 0.2 0 0 0EaB (%)259 265 110 495YM (MPa) UTS (MPa)10.55 4.91 8.00 4.482.36 2.20 N/A 1.06Table 7 – Part performance: Comparison between HSS, LS and IM2.

Thermal PropertiesFigure 13 shows the thermal scan of TPE210-S mixed with 0.2% by weight Cab-O- Sil compared to the thermal scan of virgin TPE210-S. The two curves are identical in shape, with physical transitions occurring at constant temperatures Tg (105°C), Tm (140°C) and Tc (50°C). Minor distortions are observed which may be due to the effect of impurity during mixing or irregular sample composition in the pan.Tg2Tm2Tg1Tm1Virgin TPE210-SVirgin TPE210-S + 0.2% Cab-O-SilTc1Tc2Figure 13 – DSC scan of TPE210-S mixed with 0.2% Cab-O-Sil and virgin TPE210-SThe mix powder curve (outer) have significantly shifted in the direction of heat flow at a value of around 5 mW from the virgin curve (inner). Both melt peak (Tm2) and crystallisation peak (Tc2) appear sharper in the mix powder curve compared to Tm1 and Tc1.

Both melt peak (Tm2) and crystallisation peak (Tc2) appear sharper in the mix powder curve compared to Tm1 and Tc1. This suggests the addition of flow agent has increased the amount of energy required to change the material from solid to liquid i.e heat of fusion, ΔHm and the amount of energy released during crystallisation, ΔHc. As both transitions take place at the constant temperatures of 140°C and 50°C, consequently higher thermal energy is involved in both phase changes. The presence of silica atoms in between the TPE atoms may have an effect on the molecules arrangement, making them closely packed together. This subsequently strengthens the intermolecular bonds thus more energy is needed to break them. It is also observed that the range of Tg in both curves differed slightly, with ΔTg2 higher than ΔTg1, by approximately 5°C. The process window is widened thus providing advantage to the sintering process at lower temperature range.

The process window is widened thus providing advantage to the sintering process at lower temperature range. In conclusion, the addition of flow agent in powder does not alter the original physical properties of powder, instead enhance its thermal properties.3. Surface RoughnessThe effect of adding flow agent into the powder mix can be observed through the following graph of average roughness variation in Figure 14.50.00) n o r c i mParameter 1Parameter 2(a R40.00, s s e n h g u o R e g a r e v A30.0036.3535.00 33.0231.99 30.3725.6220.000.00.20.51.0Flow agent percentage by weight (%)Figure 14 – Average roughness values against flow agent additionThe addition of flow agent generally reduces the surface roughness by promoting better flow between the particles during sintering.

However the values of Ra are still relatively high compared to parts produced by other manufacturing technologies and materials.ConclusionsThe experiments have shown that High Speed Sintering process produces more superior parts compared to equivalent Laser Sintered parts. The addition of fumed silica decreases part’s surface roughness. However further addition of fumed silica results in a decrease of Elongation at Break, evident increase of Young’s Modulus and mixed results in Ultimate Tensile Strength.References1. Gibson, I, et al. Additive manufacturing technologies: rapid prototyping to direct digital manufacturing. New York : Springer, 2010.2. Hopkinson, N, Hague, R and Dickens, P. Rapid Manufacturing: an industrial revolution for the digital age. UK : Wiley, 2006.3. Wohlers, T. Wohlers report 2007: state of the industry: annual worldwide progress report. USA: Wohlers Associates,Inc, 2007.4.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEFFECT OF SPARSE-BUILD INTERNAL STRUCTURE ON PERFORMANCE OF FUSED DEPOSITION MODELING TOOLS UNDER PRESSURES. Meng, L. Mason, G. Taylor, X. Wang, M. C. Leu, and K. Chandrashekhara Department of Mechanical and Aerospace Engineering Missouri University of Science and Technology, Rolla, MO 65409M. Matlack and J. Castle The Boeing Company, St Louis, MO, 63134AbstractTwo different approaches to design a sparse-build tool for fabrication by the fused deposition modeling (FDM) process are compared. One approach uses a 2D lattice structure and the other approach is inspired by topology optimization. Ultem 9085 is used as the material, and the amount of material used to build the tool is kept constant to ensure a fair comparison.

Ultem 9085 is used as the material, and the amount of material used to build the tool is kept constant to ensure a fair comparison. A solid tool is also included in the comparison. The performance of the tool under uniform pressure is simulated using finite element analysis (FEA) and the accuracy of the FEA results is verified by comparing them with experimentally measured data for a similar tool. The build material, support material, build time, maximum displacement, and maximum von Mises stress are compared for the three build approaches, with an emphasis on the pros and cons of each sparse- build tool with regards to performance under uniform pressure and fabrication by FDM.IntroductionShortening the product development cycle time is a key factor in ensuring that a company remains ahead of its competitors.

Additive manufacturing is a method of rapidly and efficiently fabricating a physical part based upon a design model, and it has been in use throughout various industries over the past three decades [1]. One of the major advantages of additive manufacturing is that the build time for product development is shortened significantly. Furthermore, additive manufacturing enables light-weight parts to be manufactured by changing the internal structure of the model from solid to sparse [2].Additive manufacturing is a process where layers of material are formed and added to create three-dimensional parts. One of the processes for additive manufacturing is Fused Deposition Modeling (FDM). In this process, the material filament is heated just above the melting point and extruded through a nozzle to form the layers. Materials for FDM include: thermoplastics such as ABS, Ultem 9085, Ultem 1010, polycarbonate, and many others [3].

Materials for FDM include: thermoplastics such as ABS, Ultem 9085, Ultem 1010, polycarbonate, and many others [3]. The FDM machines used to fabricate these parts offer a variety of customizable build parameters including: build direction, raster angle, raster width, raster air gap, wall/contour thickness, and cap thickness.Topology optimization has become an important research area in structural design. It utilizes a mathematical approach that optimizes material distribution within a desired design space. The optimization is based on an objective function with constraints in response to a set of loads and boundary conditions [4]. The most common objectives used in topology optimization2337are the minimization of compliance or mass. Additional constraints, such as maximum allowable stress or displacement, can be added to direct the optimization process and produce even more desirable results.

Additional constraints, such as maximum allowable stress or displacement, can be added to direct the optimization process and produce even more desirable results. The topology optimization software used for this study, solidThinking Inspire 2015, utilizes the SIMP method along with a mass/volume filter. Topology optimization works with an existing CAD model to help design structural parts and has the ability to reduce the cost, development time, material usage, and weight of a product.Finite element analysis is widely used to model additive manufacturing processes. However, literature data on modeling and simulation of FDM parts is still limited. Zhang and Chou [5] constructed a thermo-mechanical finite element model to simulate stress accumulation and the distortion of parts during the filament deposition process. Miquel et al. [6] built an effective finite element model which was based on tension tests performed on FDM parts.

Miquel et al. [6] built an effective finite element model which was based on tension tests performed on FDM parts. This verified finite element model can predict proper material orientations under varying loading conditions. Mostatfa et al. [7] used ANSYS to simulate ABS-iron composite FDM manufacturing process including various build parameters. Li et al. [8] developed a finite element model to simulate the vacuum assisted resin transfer molding process for composites with molds manufactured with the FDM process. The FDM part was optimized based on this verified finite element model. Finite element analysis has shown to be an effective method to study and optimize FDM parts.In the current study, a tool with geometry based on composite molding tools is investigated for the effects of replacing the interior of the tool with sparse-build internal structures.

Two different sparse-build internal structures are used: a 2D lattice structure called sparse double dense (SDD) and a topology optimization inspired (TOI) design. To ensure the comparability, the material volumes of the SDD and TOI tools are kept the same. A tool with a solid internal structure is used as a reference in the comparison. The performance of the tool under uniform pressure is simulated using FEA. The accuracy of the FEA results is verified by comparing them with experimentally measured data for a similar tool. The build material, support material, build time, maximum displacement, and maximum von Mises stress are compared for the three build approaches.Tool Geometry and Internal Structure GenerationTool GeometryThe geometry of the tool used for this study is shown in Figure 1a. It is a square plate with a curved top surface. An illustration of the uniform pressure loading condition is shown in Figure 1b.

It is a square plate with a curved top surface. An illustration of the uniform pressure loading condition is shown in Figure 1b. The model has a uniform pressure load of 0.6895 MPa on all surfaces except the bottom surface, where a rigid flat plate provides the boundary conditions. This geometry and loading condition are based off of common composite molding tools.2338(a) (b) Figure 1: (a) Tool geometry and (b) loading conditionSparse Double Dense DesignSparse double dense (SDD) is a light weight internal structure commonly used in FDM tooling and manufacturing applications. This internal structure has the advantages of a reduced amount of build material, shorter build time, and higher strength-to-mass ratio compared to a solid internal structure. The SDD parameters used in this study are shown in Figure 2. The raster air gap and raster angle are kept at the default values of 2.032 mm and 45 degrees, respectively.

The SDD parameters used in this study are shown in Figure 2. The raster air gap and raster angle are kept at the default values of 2.032 mm and 45 degrees, respectively. A single contour, or layer outline, is chosen instead of multiple contours to minimize the effect of the contour on the study of the internal structure. Since the extrusion tip, T16, is capable of widths from 0.4064 mm to 0.7620 mm, a raster width closest to the midpoint is chosen at 0.6096 mm.Figure 2: Sparse double dense cross section and parameters2339Topology Optimization Inspired DesignThe topology optimization design is conducted using Altair HyperWorks solidThinking Inspire 2015. This software can perform optimizations using two different objectives: minimize mass or maximize stiffness. For the minimize-mass objective, there is an optional minimum safety factor constraint.

This software can perform optimizations using two different objectives: minimize mass or maximize stiffness. For the minimize-mass objective, there is an optional minimum safety factor constraint. For the maximize-stiffness objective, there is an optional mass target constraint where the user can specify the target mass as a percentage of the design space or as a total mass value. In this study, the maximize-stiffness objective is used and the mass target is specified as a total mass value equaling that of the SDD tool.Topology optimization has a mesh dependency, meaning that different mesh densities affect the results. Due to this dependency, denser meshes create more intricate topology optimization results.

Due to this dependency, denser meshes create more intricate topology optimization results. As shown in Figure 3, the topology optimization study is completed with four different mesh densities governed by the average tetrahedral element size: low (10.16 mm), mid-low (5.08 mm), mid-high (2.54 mm), and high (1.15 mm).(a)(b)(c)(d)Figure 3: Cross sections of topology optimization results with four different mesh densities governed by the average tetrahedral element size: (a) low (10.16 mm), (b) mid-low (5.08 mm), (c) mid-high (2.54 mm), and (d) high (1.15 mm)As the topology optimization results get more and more refined with increasing mesh density and number of elements, the run time for the optimization increases. The high mesh density optimization run takes over 24 hours to complete with an element size of 1.15 mm. In this study, run times longer than this are not feasible.

The high mesh density optimization run takes over 24 hours to complete with an element size of 1.15 mm. In this study, run times longer than this are not feasible. The results shown in Figure 3 show the tendency for the topology optimization structure to split up into an increasing number of 0 and 90 degree walls as the mesh density increases. These walls can be connected together in such a way as to form nesting squares, or stiffeners, with a solid square in the middle as the smallest stiffener. Following this convention, the low mesh density (10.16mm) can be approximated by two square stiffeners, the mid-low (5.08mm) by three, the mid-high (2.54mm) by four, and the high (1.15mm) by five. It is logical that further increasing the mesh density for the topology optimization would result in an increasing number of these stiffeners.

It is logical that further increasing the mesh density for the topology optimization would result in an increasing number of these stiffeners. Inspired by this pattern of an increasing number of stiffeners, designs are created with 5, 10, 15, and 20 stiffeners, as shown in Figure 4.5 stiffeners10 stiffeners15 stiffeners20 stiffenersFigure 4: Topology optimization inspired designs2340Results and DiscussionAn FEA study is run on the TOI tools shown in Figure 4 and the performance of the four tools are compared. The tool with 10 stiffeners has the lowest maximum von Mises stress and the lowest maximum displacement. To substantiate this finding and populate more data, TOI tools with 7, 8, 9, 11, 12, 13, and 17 stiffeners are modeled and included in the FEA study. The maximum von Mises stress and maximum resultant displacement results for the TOI tools are shown in Figure 5 and Figure 6, respectively.

The maximum von Mises stress and maximum resultant displacement results for the TOI tools are shown in Figure 5 and Figure 6, respectively. As the number of stiffeners increases, the maximum von Mises stress and the maximum displacement both decrease until they reach 10 stiffeners; at which point the maximum von Mises stress remains relatively constant and the maximum displacement starts increasing. Since minimizing the maximum displacement and maximum von Mises stress is desirable for this tool, the tool with 10 stiffeners is the logical choice based on the results.

The dark blue indicates the minimum displacement and the dark red the maximum displacement. The maximum resultant displacement occurs at the curved edges between the top surface and the front and back surfaces. These edges deform inward towards the middle of the tool during the simulation. The maximum resultant displacement of the SDD (0.0118 mm) and TOI (0.0173 mm) tools are an order of magnitude larger than the solid tool (0.00148 mm). The resultant displacements along the sides of the solid and SDD tools are fairly constant along the length. This suggests that the SDD tool deforms in a manner very similar to the solid tool. The TOI tool, however, has a distinctly different deformation pattern. The areas of minimum resultant displacement on the top surface of the TOI tool run along the diagonals of the square instead of the side edges like the other tools.

The z-axis displacement contour plots are shown in Figure 8. The dark blue indicates the maximum negative (downwards) displacement and the dark red the maximum positive (upwards) displacement. Both sparse-build tools show their internal support structure clearly in these diagrams. The light blue shows where the shell is supported and the dark blue shows where depressions in the shell have formed, called dimples. The magnitude of the dimpling can be described as the difference between the maximum displacement of an unsupported area and the minimum displacement of the supported areas that are directly adjacent. The dimpling magnitude can be roughly estimated from Figure 8 to be 0.001 mm for the SDD tool and 0.002 mm for the TOI tool.

The dimpling magnitude can be roughly estimated from Figure 8 to be 0.001 mm for the SDD tool and 0.002 mm for the TOI tool. The SDD tool has a smaller maximum displacement and less significant dimpling than the TOI tool.2342Solid (mm)SDD (mm)TOI (mm)1.92E-04 7.00E-05 5.00E-05 -2.90E-04 -4.20E-04 -5.40E-04 -9.00E-04 -1.03E-03 -1.27E-031.28E-03 7.25E-04 1.75E-04 -3.76E-04 -2.03E-03 -3.13E-03 -3.68E-03 -4.23E-03 -5.33E-032.88E-03 2.16E-03 1.44E-03 7.17E-04 -7.22E-04 -2.88E-03 -3.60E-03 -4.32E-03 -5.76E-03Figure 8: Z-axis displacement contour plotsThe amount of build material, support material, and build time for each tool are estimated using the Insight 10.2 and Control Center 10.2 software. This software is created by Stratasys specifically for their machines.

This software is created by Stratasys specifically for their machines. Insight allows the user to process individual STL files by slicing them into layers, placing support material, creating tool paths for each layer, and then sending a CMB file with that information to the Control Center. The Control Center then allows the user to configure the placement of one or more CMB files on the build platform, estimates the amount of build material, support material, and time required to build the parts, and sends the information to the machine. Table 1 shows the estimated tool fabrication requirements as well as the maximum resultant displacement and maximum von Mises stress from the FEA study. Figure 9 shows a side by side comparison of the results from Table 1.The SDD and TOI tools require the same amount of build material, 64% of the solid tool.

Figure 9 shows a side by side comparison of the results from Table 1.The SDD and TOI tools require the same amount of build material, 64% of the solid tool. The solid and SDD tools require no support material besides the base that each part is printed on, but the TOI tool requires 8.35 cm3 of support material inside the tool to support the top shell in between the stiffeners. This poses a problem as the support material inside the TOI tool cannot be removed and would affect the performance of the tool.

This poses a problem as the support material inside the TOI tool cannot be removed and would affect the performance of the tool. Since the properties of the support material are not given, it is not possible to include this in the FEA study.Table 1: Estimated tool fabrication requirements and FEA resultsBuild Material (cm3)Support Material (cm3)Build Time (min)Maximum Resultant Displacement (mm)Maximum von Mises Stress (MPa)Solid SDD TOI44.41 28.35 28.351.64 1.64 9.9948 35 1110.00148 0.0118 0.01730.75 4.51 4.292343The SDD tool is expected to have a shorter build time than the solid tool and the results are consistent with this expectation. As far as fabrication goes, the only differences between solid and SDD are the raster air gap (which is zero for solid parts) and the number of extrusions per layer.

As far as fabrication goes, the only differences between solid and SDD are the raster air gap (which is zero for solid parts) and the number of extrusions per layer. Increasing the raster air gap not only decreases the amount of material, it also decreases the number of times the nozzle has to go back and forth to fill the layers, thus decreasing the build time. The TOI tool, on the other hand, has a complicated internal geometry compared to the other tools. For most of the layers, each solid area in the cross section requires 2-4 extrusions to create, and each area filled with support material requires at least 2 extrusions. That takes a minimum of 62 extrusions for the most complicated layers of the TOI tool, compared to a maximum of 3 extrusions for SDD and 2 for solid. Every time the nozzle switches between the build material and support material, it pauses to change the temperature of the heating element that melts the material.

Every time the nozzle switches between the build material and support material, it pauses to change the temperature of the heating element that melts the material. These pauses add a lot of time to the total build time, making the TOI tool take over three times as long to build as the SDD tool.

However, if the TOI tool is built without support material, the build time would be about the same as the solid tool, which suggests that transitioning from one material to the other and back again for each layer increased the build time by over 130%.) 3m c (l a i r e t a M d l i u B45 40 35 30 25 20 15 10 5 0) 3m c (l a i r e t a Mt r o p p u S10 9 8 7 6 5 4 3 2 1 0120100) n i m80(e m T d l i u Bi6040200Solid SDD TOISolid SDD TOISolid SDD TOIt n a t l u s e R m u m i x a M)m m(t n e m e c a l p s i D0.018 0.016 0.014 0.012 0.010 0.008 0.006 0.004 0.002 0.000s e s iM n o v m u m i x a M) a P M(s s e r t S5.0 4.5 4.0 3.5 3.0 2.5 2.0 1.5 1.0 0.5 0.0Solid SDD TOISolid SDD TOIFigure 9: Results comparison2344The compressive yield strength of Ultem 9085 is 87-100 MPa, depending on the orientation of the load with respect to the print direction.

Using the lower bound of the compressive yield strength range, the tools have a factor of safety of 116 for the solid tool, 19 for the SDD tool, and 20 for the TOI tool. All three tools can safely withstand the 0.6895 MPa pressure load.Experimental VerificationSince the uniform pressure loading condition cannot be experimentally tested for this study, an FEA study is performed to simulate the uniform pressure results. To ensure the accuracy of the results for the original tool, an FEA study is run on a verification tool and the results are compared to a compression test experiment. The machine used for the compression testing is an INSTRON 5985. The verification tool is a modified version of the original tool geometry in this study.

The machine used for the compression testing is an INSTRON 5985. The verification tool is a modified version of the original tool geometry in this study. As shown in Figure 10, the verification tool has a flat top to create a uniform distributed load on the top of the tool during the compression test.Figure 10: Verification tool geometryThe SDD internal structure is applied and a topology optimization is performed on the verification tool in the same manner as the original tool. The topology optimization result is shown in Figure 11, which is then modified to have a single flat vertical surface on either side instead of the triangular mesh surface.Figure 11: Topology optimization result for the verification tool2345The verification tools are fabricated using a Stratasys Fortus 400mc machine with Ultem 9085 as the print media. Five samples each of the solid, SDD, and topology optimized verification tools are fabricated and then compression tested.

Five samples each of the solid, SDD, and topology optimized verification tools are fabricated and then compression tested. The experimental and FEA results are shown in Figure 12. The difference between the FEA and experimental results are 0.45% for solid, 7.66% for SDD, and 6.16% for topology optimization (TO). The results are in good agreement; therefore the accuracy of the FEA results is verified.0.30t n e m e c a l p s i D m u m i x a M)m m(0.250.200.150.100.05FEAExperimental0.00SDD Figure 12: Verification tool resultsSolidTOConclusionA tool, with the geometry and loading condition based off of composite molding tools, is investigated for the effects of replacing the interior of the tool with two sparse-build internal structures: sparse double dense (SDD) and a topology optimization inspired (TOI) design.

The performance of the tool under uniform pressure is simulated using FEA.The maximum resultant displacements of the SDD (0.0118 mm) and TOI (0.0173 mm) tools are an order of magnitude larger than that of the solid tool (0.00148 mm). The build material for the SDD and TOI tools are kept constant at 64% of the solid tool. The TOI tool requires interior support material, which increases the build time by over 130% and cannot be removed. The tools have a factor of safety of 116 for the solid tool, 19 for the SDD tool, and 20 for the TOI tool. Thus all three tools can safely withstand the 0.6895 MPa pressure load.For the composite molding application of this tool, the outer surface displacement has the greatest impact on the quality of the composite product. It stands to reason that the SDD tool would produce a higher quality product than the TOI tool as it has a lower maximum displacement, less significant dimpling, and a similar displacement pattern to the solid tool.

The SDD tool is also the logical choice when additive manufacturing is taken into consideration since it has the shortest build time and requires no extra support material. The TOI tool has a slightly lower maximum von Mises stress than the SDD tool, but this advantage can be considered negligible due to the excessively large factor of safety. Therefore, for this study the SDD internal structure is more desirable than the TOI internal structure.2346AcknowledgementsThis research is sponsored by the Industrial Consortium of the Center for Aerospace Manufacturing Technologies (CAMT) at Missouri University of Science and Technology. Special thanks to Michael Hayes of Boeing and Timothy Schniepp and Chris Holshouser of Stratasys for their valuable contributions to the project.References[1] C. K. Chua, K. F. Leong, and C. S. Lim, Rapid Prototyping: Principles and Applications, vol. 1, 2nd ed. Singapore, Singapore: World Scientific Publishing Co Pte, 2003, pp. 124–129.[2] O.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEFFECT OF SPECIMEN SURFACE AREA SIZE ON FATIGUE STRENGTH OF ADDITIVELY MANUFACTURED TI-6AL-4V PARTSJonathan Pegues1, Michael Roach2, R. Scott Williamson2, Nima Shamsaei1,*1Laboratory for Fatigue & Additive Manufacturing Excellence, Department of Mechanical Engineering, Auburn University, Auburn, AL 368492Department of Biomedical Materials Science, University of Mississippi Medical Center, Jackson, MS 39216Corresponding author: shamsaei@auburn.eduAbstractAs additive manufacturing becomes an increasingly popular method for advanced manufacturing of components, there are many questions that need to be answered before these parts can be implemented for structural purposes.

One of the most common concerns with additively manufactured parts is the reliability when subjected to cyclic loadings which has been shown to be highly sensitive to defects such as pores and lack of fusion between layers. It stands to reason that larger parts will inherently have more defects than smaller parts which may result in some sensitivity to surface area differences between these parts. In this research, Ti-6Al-4V specimens with various sizes were produced via a laser-based powder bed fusion method. Uniaxial fatigue tests based on ASTM standards were conducted to generate fatigue-life curves for comparison. Fractography on the fractured specimens was performed to distinguish failure mechanisms between specimen sets with different sizes.Keywords: Fatigue; Additive manufacturing; Surface effects; Microstructure; Titanium alloysIntroductionAdditive manufacturing (AM) of components has become an increasingly popular advanced manufacturing process in many industries.

In the biomedical industry the AM process offers customized implants that can be directly tailored to an individual. Titanium alloys such as Ti-6Al-4V are widely used in biomedical applications due to their excellent biocompatibility [1 - 4]. Traditionally these implants have been produced via subtractive manufacturing from wrought materials in which extensive research into the durability under various conditions has been conducted [4 – 10]. One of the biggest challenges AM produced materials must overcome before being fully implemented into structural and biomedical applications is the uncertainty of the fatigue behavior compared to traditional wrought alloys [11]. Research has shown that the fatigue life of AM parts shows a reduction in fatigue limits which is even greater for as-built parts that don’t undergo post-process machining to remove the rough surface [12 – 14].

While much improvement in the additive process has been achieved in recent years, there is still a need to better understand the fatigue behavior of AM parts before they are fully implemented into structural applications.122AM parts have shown high sensitivity to defects inherent to the AM process such as voids and lack of fusion (LoF) between layers [14 – 16]. LoF between layers typically results in seemingly fully dense parts, however, these “pancake like” voids contain little volume yet create high stress concentrations along the edges. As such, the LoF areas are prime crack initiation sites which contribute to the reduced the fatigue life observed in AM parts. Additionally, for as-built parts, the effect of surface area may prove to be detrimental to the fatigue limit due to roughness induced crack initiation along the surface.

Additionally, for as-built parts, the effect of surface area may prove to be detrimental to the fatigue limit due to roughness induced crack initiation along the surface. Parts that contain similar volume with varying surface areas may result in noticeably different fatigue lives for similar stress amplitudes and thus lower fatigue limits. This research specifically addresses the issue of increasing surface area for AM parts produced via a laser-based powder bed fusion (L-PBF) method through the use of several microscopic techniques.Materials and MethodsThree sets of specimens were designed to produce increasing gage surface area while maintaining a constant gage volume of approximately 123 mm3. The chosen specimen geometries are shown in Figure 1 along with the calculated surface area for each specimen type. Specimens were fabricated using a powder bed based EOS M290 with a maximum laser power of 400W.

Specimens were fabricated using a powder bed based EOS M290 with a maximum laser power of 400W. The spherical Ti-6Al-4V alloy powder was supplied with a powder size greater than 15 m and less than 45 m. All specimens were built at an angle of 45° from the titanium substrate. Before removal, the specimens were stress relieved to alleviate any residual stress as result of the extreme cooling rates associated with the L-PBF AM process. The stress relief was conducted at 700° C for 1 hr under an Ar environment and allowed to air cool after which the samples were carefully removed from the substrate.Figure 1: Sample dimensions. All measurements are in mm.123A representative sample was sectioned to determine the as-built microstructure. The microstructure was revealed using Kroll’s reagent to etch the polished surface which was then observed using an optical microscope.

The microstructure was revealed using Kroll’s reagent to etch the polished surface which was then observed using an optical microscope. Further microstructural analysis was performed using scanning electron microscopy (SEM) and electron backscatter diffraction (EBSD) analysis to determine the present phases and grain size. Fatigue tests were conducted on an MTS testing frame with a load capacity of 100 kN. An extensometer was used to record the stabilized hysteresis response for each test then removed so that the samples could be run until failure. Three stress levels were selected to produce a range of fatigue lives from 104 cycles and greater than 106 cycles. At least two and up to four samples were tested at each stress level for all sample geometries to construct a stress-life (S-N) curve for comparison between the sample types. All tests were run until failure except for tests that exceeded 106 cycles which were removed and considered to be runouts.

All tests were run until failure except for tests that exceeded 106 cycles which were removed and considered to be runouts. Fractured surfaces were subsequently removed and observed by SEM to identify failure mechanisms related to each specimen.Experimental ResultsMicrostructureThe microstructure consisted of mostly acicular α’ grains formed as a result of diffusionless phase transformation from the  phase with a small distribution of lath like α phase. This   α’ phase transformation is a result of the extraordinarily high cooling rates associated with the L-PBF process and results in a fully α’ + α phase distribution [17, 18]. The microstructure developed in an epitaxial manner resulting in elongated prior  grains oriented parallel to the build direction as shown in Figure 2.

The microstructure developed in an epitaxial manner resulting in elongated prior  grains oriented parallel to the build direction as shown in Figure 2. The top surface in relation to the build plate had a visually smoother surface compared to the bottom surface as indicated in Figure 2.Figure 2: Optical image taken at 50x in the transverse direction. The left side shows upper most portion of the build while the right side shows the bottom most portion.124The microstructure and phase distribution were revealed using EBSD. A representative sample from the transverse direction is shown in Figure 3. The α and α’ phases are both hexagonally close packed (HCP) phases and are therefore indistinguishable from each other in the phase map as shown in Figure 3(b). However, the acicular needle like α’ phase is visually distinguishable from the lath like α phase within the prior  grains which is shown by the dotted line in Figure 3(a).

However, the acicular needle like α’ phase is visually distinguishable from the lath like α phase within the prior  grains which is shown by the dotted line in Figure 3(a). The phase map does show that the  phase was completely transformed to the α or α’ phases upon cooling. Grain size analysis showed similar grain size and shape for the α and α’ phases in both the longitudinal and transverse direction which is given in Figure 3(c) and (d) respectively. The average width of the α’ grains was approximately 1.5 m with an average aspect ratio of 0.55.Figure 3: Microstructure of the as-built microstructure showing a) orientation b) phase distribution c) grain size and d) aspect ratio of the grains.Fatigue BehaviorThe S-N behavior of wrought Ti-6Al-4V has been well studied and has been shown to have a fatigue limit in the range of 550 - 750 MPa [19].

For additively manufactured Ti-6Al-4V, however, the fatigue limit has been shown to be significantly reduced, especially for the as-built condition [14]. Fatigue tests were started at a stress amplitude of 450 MPa to establish a sufficient set of stress levels to produce fatigue lives in the desired range. The stress amplitude of 450 MPa produced relatively low cycle fatigue lives (< 20,000 cycles) which is in stark contrast to the fatigue limit of wrought Ti-6Al-4V alloys. It was found that a stress level of 100 MPa was sufficient in producing fatigue lives greater than 106 cycles which for this investigation was considered to be approaching the fatigue limit. From these results the stress levels selected to125construct S-N curves for the given geometries were 300 MPa, 200 MPa, and 100 MPa.

From these results the stress levels selected to125construct S-N curves for the given geometries were 300 MPa, 200 MPa, and 100 MPa. At least one test for each sample set was conducted at 450 MPa, however, they were not included in the cyclic deformation analysis due to the inconsistent deformation behavior for low cycle fatigue under load controlled conditions.Cyclic data from the fatigue tests showed that all stress levels resulted in fully elastic deformation as would be expected for Ti-6Al-4V under such low stress amplitudes. Figure 4 shows the hysteresis loops for samples #1 and #3 and show fully elastic deformation for both samples at 300 MPa.

Figure 4 shows the hysteresis loops for samples #1 and #3 and show fully elastic deformation for both samples at 300 MPa. All sample geometries resulted in similar cyclic behavior for each stress amplitude and show almost no macroscopic plastic deformation.Figure 4: Hysteresis loops for a) sample set #1 and b) sample set #3.The S-N curve for each sample geometry is given in Figure 5 and shows a clear separation in the stress life of the respective geometries. The arrow in Figure 5 represents tests that were removed once sufficiently past 106 cycles without failure. Sample geometries with increased surface area showed a general reduction in fatigue lives at all stress levels at which failure occurred. For the intermediate fatigue lives at a stress amplitude of 200 MPa there is some overlap in the fatigue life ranges between each sample set.

For the intermediate fatigue lives at a stress amplitude of 200 MPa there is some overlap in the fatigue life ranges between each sample set. However, it can be reasoned that there is an observable reduction of fatigue life for this stress level when considering the average fatigue life for each geometry. For stress amplitudes over 200 MPa the separation in fatigue lives is much clearer for each geometry without overlap between the respective fatigue life ranges. Additionally, the low cycle fatigue lives at 450 MPa are shown in Figure 5.

Additionally, the low cycle fatigue lives at 450 MPa are shown in Figure 5. While these tests contain high uncertainty due to the low cycle fatigue behavior for tests in load control conditions, there is an obvious separation in fatigue life at this stress level as well.126Figure 5: Stress-life curves for each sample geometry.FractographyFractography showed that the fracture surfaces were somewhat similar across all sample geometries in that there were three distinct crack growth zones associated with the fracture surfaces. These zones are shown in Figure 6 where the arrows indicate the crack growth directions.Figure 6: Representative crack growth zones for tested samples.127Interestingly, all cracks initiated on the side that was facing the build plate during fabrication indicating that AM parts in the as-built condition are highly sensitive to build orientation and specifically to any surfaces oriented towards the build plate.

Cracks initiated at the surface and almost exclusively at a point where two powder particles were melted together as shown in Figure 7. The arrows in Figure 7 indicate the initiation sites where two powder particles are melted together creating a micro-notch in the already highly rough surface. A tear ridge is formed at the right most initiation site and extends upwards which indicates a change in grain orientations across the tear ridge, possibly a prior  boundary. The crack initially propagates in stage I crack growth mode along crystallographic planes as indicated by the highlighted (red) facets. These facets resemble the martensitic α’ needles that were revealed during the microstructure analysis.

These facets resemble the martensitic α’ needles that were revealed during the microstructure analysis. This initial type of crack growth leads to a relatively tortuous crack path as the crack tends to change directions when encountering discontinuities such as grain boundaries.Figure 7: Crack initiation site at joining powder particles (arrows) showing stage one crack growth exposing α' facets (red).As the crack propagated, a mixture of transgranular and intergranular crack growth along with large secondary cracking and fatigue striations perpendicular to the crack growth direction were observed. Figure 8 gives a representation of these characteristics which were representative of the secondary crack growth zone shown in red in Figure 6. The exposed grain in the center of Figure 8 shows quasi-cleavage characteristics in which an apparent crystallographic plane ruptured while leaving tear dimples across the surface.

The exposed grain in the center of Figure 8 shows quasi-cleavage characteristics in which an apparent crystallographic plane ruptured while leaving tear dimples across the surface. To the right of the grain in Figure 8, fatigue striations are visible showing the crack front was propagating towards the top left of the image. These features indicate that this secondary zone of crack propagation is a mixture of static failure modes such as cleavage and ductile rupture as well as typical stage II fatigue crack growth with stable striation marks. This combination of failure modes leads to an even more tortuous crack path than the first crack growth zone which is readily observable in Figure 6.

This combination of failure modes leads to an even more tortuous crack path than the first crack growth zone which is readily observable in Figure 6. The combination of static and128cyclic failure modes could explain the drastically reduced fatigue limit for the L-PBF produced Ti-6AL-4V parts.Figure 8: SEM image of the second crack growth zone associated with the AM produced parts.The final crack growth zone is characterized by severe overload as a result of the decreasing load bearing cross-section from the extending crack. At this stage the part fails in a ductile overload manner. Figure 9 gives a representation of the fracture surface at this stage.

At this stage the part fails in a ductile overload manner. Figure 9 gives a representation of the fracture surface at this stage. The dimple rupture is indicative of micro-void coalescence until the remaining part fails in shear mode as shown by the steep surface along the back ridge of the specimen in Figure 6.Figure 9: Dimple rupture surface for the third crack growth zone indicating ductile overload failure.129DiscussionIt has been well established that increasing surface roughness is detrimental to the fatigue behavior of most materials [20]. Interestingly, this research has shown that, for additive manufactured parts in the as-built condition with high surface roughness, the effect is compounded by increasing surface area of a given part. The S-N curves presented in Figure 5 show the obvious trend of decreasing fatigue lives for increasing surface area, all of which maintained a similar gage volume size of 123 mm3.

The S-N curves presented in Figure 5 show the obvious trend of decreasing fatigue lives for increasing surface area, all of which maintained a similar gage volume size of 123 mm3. The detrimental effect of increasing surface area is, in part, a result of increasing crack initiation sites. Figure 10 shows a representation of the crack initiation sites associated with each sample geometry for a stress amplitude of 200 MPa. While sample set #1 and #2 showed similar numbers of crack initiation sites, sample set #3 showed significantly more initiation sites and shorter lengths until cracks began the coalescence process.

While sample set #1 and #2 showed similar numbers of crack initiation sites, sample set #3 showed significantly more initiation sites and shorter lengths until cracks began the coalescence process. With increasing surface area, the likelihood that multiple crack initiation sites will occur near one another is increased leading directly to earlier crack coalescence and shorter fatigue lives.Figure 10: Crack initiation sites for a) sample set #1 b) sample set #2 and c) sample set #3.130Additionally, the first two crack growth zones were found to be a larger for the smaller surface area samples. This is shown in Figure 11 for samples #1 and #3 which had the smallest and largest surface areas respectively. This also appears to be a result of the increased likelihood of multiple cracks initiating near each other and crack coalescence occurring much earlier in the fatigue life.

This also appears to be a result of the increased likelihood of multiple cracks initiating near each other and crack coalescence occurring much earlier in the fatigue life. The stable crack growth zone (blue) in Figure 11 was determined to cover approximately 30% of the fracture surface for sample set #1 compared to just 16% for sample set #3. Additionally, the quasi-stable crack growth zone (red) in Figure 11 covers approximately 55% for sample set #1 and 40% for sample set #3. The final ductile overload zone (white) is also significantly more torturous for sample set #3.Figure 11: Crack growth zones for a) sample set #1 and b) sample set #3SummaryIncreasing surface area for a given volume size was found to be detrimental to the fatigue life of AM L-PBF produced Ti-6Al-4V parts. This is, in part, a result of the increased likelihood that multiple cracks will initiate in close proximity to each other.

This is, in part, a result of the increased likelihood that multiple cracks will initiate in close proximity to each other. The fatigue crack growth zones for the increased surface area samples were found to be comparatively smaller than the samples with smaller surface areas. Additionally, the second crack growth zones were found to contain both static and cyclic failure modes which could explain why the fatigue endurance limit shows such a drastic reduction when compared to wrought Ti-6Al-4V alloys. These results indicate that geometry design for AM produced parts could improve the fatigue behavior for parts that will not undergo post-process machining by limiting the exposed as-built surface area.131AcknowledgementThis work was partially funded by the National Science Foundation under Grant No. 1657195. Support from the Naval Air Systems Command (NAVAIR) and Magee Technologies LLC is also greatly appreciated.References[1] C. N. Elias, J. H. C. Lima, R. Valiev, and M. A.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEFFCT OF SPRAY-BASED PRINTING PARAMETERS ON CEMENTITIOUS MATERIAL DISTRIBUTIONBing Lu1,2, Mingyang Li*1, Wenxin Lao1, Yiwei Weng1,2, Shunzhi Qian1,2, Ming Jen Tan1, Kah Fai Leong11Singapore Centre for 3D Printing, School of Mechanical and Aerospace Engineering, Nanyang Technological University, Singapore 639798 2School of Civil and Environmental Engineering, Nanyang Technological University, Singapore 639798AbstractIn the past decade, 3D printing is getting into more and more industry areas including building and construction. However, most 3D cementitious material printing processes are limited in horizontal printing surface. Due to the nature of building and construction industry, 3D spray cementitious material printing process was developed to apply material in vertical or even overhang surfaces.

Due to the nature of building and construction industry, 3D spray cementitious material printing process was developed to apply material in vertical or even overhang surfaces. Unlike traditional manually operated spray method in building and construction industry, 3D spray cementitious material printing process requires higher accuracy on material distribution. In this paper, the effects of four printing parameters (cementitious material flow rate, air flow rate, nozzle travel speed, nozzle standoff distance) on material distribution in 3D spray cementitious material printing process were investigated experimentally. An experimental model, which can be further used in the control of 3D spray cementitious material printing process, was then developed upon on the results.Keywords: Additive manufacturing, 3D cementitious material printing, spray, material distribution, modelling1.

Introduction3D printing, also known as additive manufacturing, is a method to sequentially deposit material layer by layer to fabricate desired objects [1]. Recently, remarkable progress of 3D printing has been seen in building and construction industry. Large computer-controlled gantry printing systems, e.g. Contour Crafting [2], Concrete Printing [3] have been successfully developed to print cementitious materials. On the other hand, the precise movement control of robotics contributed to robot-based printing systems [4] [5], which have less size limitation of printed structure compared with gantry printing system. The successful practices of these printing systems stimulate the corresponding material research. Le et al. [6] proposed several key properties for high performance printable cementitious materials, e.g. pumpability and buildability.

Le et al. [6] proposed several key properties for high performance printable cementitious materials, e.g. pumpability and buildability. Subsequent research studies [7] [8] [9] [10] further revealed that pumpability and buildability are closely related to rheological properties of the material. With these material development guidance, functional printable cementitious materials have been invented, e.g. 3D printable sustainable material [11], 3D printable fiber reinforced concrete material [12] [13].Most 3D printing in the building and construction is extrusion-based, and the shape of extrudate can be effectively controlled by the design of nozzle outlet [14], nozzle standoff height [15], printing speed [16], etc. However, there is little study of 3D cementitious material printing on the vertical surface, e.g. printing of vertical decorative pattern.

However, there is little study of 3D cementitious material printing on the vertical surface, e.g. printing of vertical decorative pattern. Different fromlimingyang@ntu.edu.sg1989printing on the horizontal surface, the vertically deposited material needs to resist the shear induced by gravity. In the building and construction area, the conventional solution is the adoption of spraying concrete materials. However, due to the lack of systematic control of spray and large rebound of high-speed sprayed material at far distance, the sprayed material on the substrate is not uniformly distributed and needs subsequent manual scraping work [17]. In contrast, 3D spray cementitious material printing process requires higher accuracy on the material distribution without human intervention. However, there is very limited study on 3D sprayable cementitious materials [18], and no research has been carried focusing on material distribution of 3D spray cementitious material printing.

However, there is very limited study on 3D sprayable cementitious materials [18], and no research has been carried focusing on material distribution of 3D spray cementitious material printing. The situation stimulates the research motivation for this paper.The material distribution of 3D spray cementitious material printing is systematically studied in this paper. Similar to extrusion-based 3D cementitious material printing, material distribution may vary with different printing parameters. The effects of four printing parameters on sprayed material distribution have been investigated, i.e. cementitious material flow rate, air flow rate, nozzle travel speed, and nozzle standoff distance. An experimental model of material distribution is hereby proposed, which can be utilized for further study of 3D spray cementitious material printing.2.

An experimental model of material distribution is hereby proposed, which can be utilized for further study of 3D spray cementitious material printing.2. Material DesignThe sprayable cementitious material used in the experiments was composed of Type-I ordinary Portland cement (OPC), Class-F fly ash, silica fume, sand and tap water. The grading curve of river sand is shown in Figure 1. Superplasticizer of 1 g and Air entraining agent (AEA) of 0.1 g was added to per liter mix. The mix proportion of the sprayable cementitious material is shown in Table 1.Figure 1 Sand gradation1990Table 1 Mix mass proportion of the sprayable cementitious materialCement1Sand / Cement Ratio 0.45Fly ash / Cement Ratio 0.5Silica fume / Cement Ratio 0.05Water / Cement Ratio 0.62Super- plasticizer1 g/LAEA0.1 g/LThe sprayable cementitious material was prepared in Hobart HL-200 mixer as per following procedure.

Weighed raw ingredients except water, AEA and superplasticizer were mixed for 180s at slow speed. The water with dissolved AEA was added afterwards and mixed with other ingredients for another 180s at slow speed. Then superplasticizer was added to the mixture, firstly mixed for 90s at slow speed and followed by another 90s mixing at high speed. Then the material was taken for the experiments.Some key fresh properties were assessed to check the pumping performance of the designed sprayable cementitious material. The flowability loss of the fresh sprayable cementitious material was traced by the flow table test, which provides flow diameter of the material after the slump cone is lifted and stroke for 25 strikes [19]. As shown in the Figure 2, there is no sudden decrease of flowability within one hour.

As shown in the Figure 2, there is no sudden decrease of flowability within one hour. The viscosity of fresh cementitious material can be described by Bingham Plastic model, and the yield stress and plastic viscosity were measured using Viskomat rotational rheometer, of which values are 115.18 Pa and 20.27 Pa∙s respectively. Pumpability test with printing delivery system showed that the material with such rheological parameters can be pumped consecutively.Figure 2 Average flow diameter with time3. Experiment DesignThe 3D spray cementitious material printing experiments were designed to be completed with robotic arm control. The robotic arm used in the experiments is ABB IRB- 1600 robotic arm with 1.45 m reach. MAI Pictor Pump was adopted for pumping material through a hose of 2.5 m in length and 25.4 mm in diameter. High pressure air was connected from lab vessel with a hose of 5 m in length and 12.7 mm in diameter.

High pressure air was connected from lab vessel with a hose of 5 m in length and 12.7 mm in diameter. The spray nozzle1991connected to the cementitious material and high pressure air hoses was attached to the end of robotic arm. The whole setup of equipment can be referred to in Figure 3.Figure 3 Experiment setup consist of robotic arm, cementitioul pump with hose and high pressure airThe nozzle’s orientation and position can be controlled by the movement of robotic arm as per designed programme. The spray nozzle is adjusted to point perpendicularly to the vertical substrate in the 3D spray cementitious material printing process. The nozzle travel path is illustrated in Figure 4. In the experiments, the nozzle travels parallel to the substrate plane at five different speeds to spray straight filaments, i.e. 50 mm/s, 100 mm/s, 150 mm/s, 200 mm/s and 250 mm/s.

In the experiments, the nozzle travels parallel to the substrate plane at five different speeds to spray straight filaments, i.e. 50 mm/s, 100 mm/s, 150 mm/s, 200 mm/s and 250 mm/s. The distances between spray nozzle and the substrate were set as three different values in different runs of experiments, i.e. 50 mm, 70 mm and 100 mm.Figure 4 Nozzle travel path with different travel speedsIn addition to the nozzle position and travel speed controlled by robotic arm, cementitious material flow rates and air flow rates were also set to have different levels in the experiments. The cementitious material flow rate is roughly proportional to the pump speeds, which were set as 600 rpm and 1200 rpm in experiments. Air flow rates can be identified by1992air inject pressure values, which were 0.5 Bar and 1.0 Bar in experiments.

Air flow rates can be identified by1992air inject pressure values, which were 0.5 Bar and 1.0 Bar in experiments. The complete experiment design table is shown in Table 2.Table 2 Experiment design tableGroupA B C D E F G H IPumping rate (rpm)600 600 600 1200 1200 1200 1200 1200 1200Air inject pressure (bar) 0.5 0.5 0.5 0.5 0.5 0.5 1.0 1.0 1.0Nozzle standoff distance (mm)50 70 100 50 70 100 50 70 100Nozzle travel speed (mm/s)50 ~ 250 50 ~ 250 50 ~ 250 50 ~ 250 50 ~ 250 50 ~ 250 50 ~ 250 50 ~ 250 50 ~ 250After completion of 3D spray cementitious material printing process, the sprayed filaments were covered with plastic sheet for 24 hours in the lab. This guarantees the sprayed filaments have enough strength to be scraped from the substrate without deformation. Three samples were cut from the centre of each sprayed filament to expose the cross section (see Figure 5). The length of each sample is 30 mm.

Three samples were cut from the centre of each sprayed filament to expose the cross section (see Figure 5). The length of each sample is 30 mm. The images of cross section were processed and analysed by MATLAB. Detailed process of image analysis can be found in Lao et al. [14].Figure 5 Exposed corss sections of three samples cut from sprayed filament4. Results and DiscussionsThickness distribution with respect to filament width is used to describe the distribution of sprayed materials. A uniform material distribution should have nearly constant thickness, i.e. the cross section approaches rectangular shape. The integration of thickness distribution is1993the cross section area, which suggests how much material is sprayed. The thickness distribution of experiment groups A to I are shown in Figure 6. It has to be pointed out that, due to very thin thickness of sprayed filaments, no data was collected for experiment group C.

It has to be pointed out that, due to very thin thickness of sprayed filaments, no data was collected for experiment group C. The integrated cross section area values are shown in Figure 7.Figure 6 Thickness Distribution (Group A to Group I)1994Figure 7 Cross section area at different travel speeds in each groupThere are some general characteristics in the material distribution. It can be found that increased pump speed can greatly increase the width and thickness of sprayed filament. However, there is no significant difference in width and thickness when air flow rate increases. The increase of nozzle travel speed decreases the thickness of sprayed filament, especially when the nozzle travel speed increases from 50 mm/s to 100 mm/s. Nevertheless, the thickness difference is not significant when the speed is higher than 100 mm/s in Group A and B; for other groups when the speed is higher than 150 mm/s the thickness difference also decreases.

Larger standoff distance can enlarge the thickness difference and also increase the width of sprayed filament. In addition, larger standoff distance leads to smaller thickness.Figure 7 shows cross section area greatly increases with higher pumping rate, especially when the nozzle travel speed is smaller than 200 mm/s. Similar to the trend of thickness distribution, there is little difference of cross section area when the nozzle travels at high speed. The phenomenon can be attributed to the discontinuity in the pumping process, which is traced from decreased density of sprayed filaments as shown in Table 3.

The phenomenon can be attributed to the discontinuity in the pumping process, which is traced from decreased density of sprayed filaments as shown in Table 3. On the other hand, either based on cross section area or density data, air flow rate does not seem to significantly contribute to material distribution.Table 3 Density of sprayed filaments (g/cm3)Travel speed (mm/s) 50 100 150 200 250A1.64 1.40 1.39 1.23 1.34B1.65 1.47 1.23 1.05 0.98C/ / / / /D1.67 1.57 1.52 1.39 1.26E1.79 1.64 1.58 1.46 1.20F1.62 1.54 1.39 1.14 0.99G1.64 1.57 1.61 1.48 1.36H1.67 1.59 1.52 1.60 1.40Table 4 and Table 5 show volume flow rate and mass flow rate of each group respectively. As nozzle travel speed increases, the volume flow rate increases while the mass flow rate remains similar when travel speed is smaller than 200 mm/s.

As nozzle travel speed increases, the volume flow rate increases while the mass flow rate remains similar when travel speed is smaller than 200 mm/s. With the same travel speed (lower than 200 mm/s), it can be found that both volume flow rate and mass flow rate1995I1.54 1.59 1.59 1.45 1.49increase proportionally as cementitious material flow rate increases. In contrast, when the nozzle travels at high speed, volume flow rate and mass flow rate increase.

Construction of Empirical ModelThere is limited study on the distribution of sprayed materials. Ginouse and Jolin proposed a second-order Gaussian distribution model to describe the mass distribution of conventional sprayed concrete materials when the spray nozzle does not move [20]. In this case, the sprayed pattern has central symmetry, and more material is gathered around the centre part. The second-order Gaussian distribution model can be expressed as Eq. (1) and Eq. (2) in polar coordination system. The origin is the nodal by central axis of spray nozzle and substrate.F  rr max  =ja 1p( ) rexp  =    r rj F max  max a 3−   0−rr maxa 2    +b 1r≤r>expr maxr max  −r rmax b 3−b 2  (1)(2)where rmax is the maximum radius of sprayed material, jp(r) is the mass flux density at radius r, and jmax is the maximum mass flux density that can be measured.

The coefficient a1, a2, a3, b1, b2, b3 in Eq. (2) can be decided by experiment data. Mass flux density refers to the mass flow rate for infinitesimal area, so mass flow rate Q (kg/s) can be expressed as Eq. (3). Mass flux density can also be estimated experimentally by sampling of sprayed material.1996Qπ= ∫ 2r max0rjp( ) r drThe experiments carried by Ginouse and Jolin showed that their proposed second-order Gaussian distribution model has good fit with experiment data when spray nozzle is placed 0.5 m and 1 m away from substrate. Refer to the good fit results of conventional spray concrete, the empirical model of 3D spray cementitious material printing may be constructed on this basis. Nevertheless, significant differences have to be taken into consideration. Firstly, the linear movement of spray nozzle make the sprayed pattern lose central symmetry.

Nevertheless, significant differences have to be taken into consideration. Firstly, the linear movement of spray nozzle make the sprayed pattern lose central symmetry. As the cross sections are vertical to the nozzle movement direction, the movement of nozzle should not affect material distribution in cross section planes. Secondly, the nozzle is placed much nearer to the substrate as per application of 3D spray printing. Therefore, the expectation of material distribution is first-order Gaussian distribution with some modifications. However, the experiment results in Figure 6 greatly deviate from Gaussian distribution. Thus the empirical model to describe 3D spray cementitious material printing cannot be constructed on the basis of Gaussian distribution.There were three steps to construct the empirical material distribution model. The first step was to check the significances of printing parameters on the maximum width of sprayed filament by statistics model.

The first step was to check the significances of printing parameters on the maximum width of sprayed filament by statistics model. The second step was considering the considering the physical deposition process and find the relationship between printing parameters and the maximum filament width. The third step was to correlate material distribution with width and average thickness data. In this step, the material distribution was described by three linear functions. Least square method was then utilized to construct the parameters of fitting functions.From experiment results, it is suspected that air flow rate does not have significant effect on material distribution. Hence, statistical checking was carried to validate this suspect. Table 6 shows p-values of printing parameters for checking their correlation with width and maximum thickness of sprayed filaments. The confidence interval is 95%.

Table 6 shows p-values of printing parameters for checking their correlation with width and maximum thickness of sprayed filaments. The confidence interval is 95%. The p-values of air inject pressure exceed 0.05 significantly in each case, which certifies that air flow rate does not have significant effect on material distribution. In contrast, all the other p-values are greatly smaller than 0.05, suggesting the other three printing parameters have significant effects on material distribution. Therefore, air flow rate will not be included in the empirical model.Table 6 p-values of printing parametersPrinting parameters Pumping rate Air inject pressure Nozzle travel speed Standoff distancep-value for width 1.23×10-6 0.58 3.09×10-6 1.02×10-4p-value for maximum thickness 0.02 0.72 1.25×10-10 0.02Considering the physics of the spray and deposition processes. When the material is sprayed out from nozzle, the spray angle can be considered as constant.

When the material is sprayed out from nozzle, the spray angle can be considered as constant. Therefore, the maximum width of filament should be linearly related to the standoff distance. At the other hand, the maximum width of filament should be proportional to the square root of filament cross section area, which is proportional to the ratio between pumping flow rate and nozzle travel speed. Denote pumping rate as α (rpm), nozzle travel speed as γ (mm/s) and standoff distance as δ (mm), the function to describe the maximum width of sprayed filament w (mm) should be in the following expression:1997(3)wmax=δ C 1α γ+C2where C1 and C2 are constants depend on the pump, nozzle and material properties.

Fitting experimental maximum filament widths into the proposed model, the proposed model described bywmax=0.145δα γ+33.9The comparison between experimental results and this fitted model is shown in Figure 8, the coefficient of determination R2 is 0.752, suggesting this model has some rationality while still needs some improvement.Figure 8 Comparison between experimental maximum width and fitted model.The next step of constructing empirical model is to correlate material distribution with width and average thickness data. By observation, the non-dimensional thickness distribution of printed filament can be roughly described by a trapezoid model( s x a b c , ,,)=    caxb −( 1x)0b a −1x b a < ≤ − 1 < ≤ x≤ ≤xb cb c1where x = w/wmax is the non-dimensional filament width coordinate, s = hwmax/(α/γ) is the non- dimensional filament thickness coordinate, a, b, and c are parameters which can be found by curve fitting.

One fitting example is shown in Figure 9. All fitted parameters of experiment groups A to I are plotted in Figure 10. It is very difficult to draw any meaningful conclusion from this distribution in current stage. While further investigation will be conducted in the future, the average value of each parameters (a = 14.5, b = 1.81, c = 9.56) were used for the prediction of material distribution in current study.1998(4)(5)(6)Figure 9 The comparison between experimental thickness and fitted three straight lines model of experiment group E with 150mm/s nozzle travel speed.Figure 10 Fitted parameter for filament thickness distribution6. Verification of Empirical ModelValidation experiments were designed to verify the proposed empirical model. The pumping rate, air inject pressure, nozzle standoff distance of validation experiments were 900 rpm, 0.75 bar, and 70 mm, respectively.

The pumping rate, air inject pressure, nozzle standoff distance of validation experiments were 900 rpm, 0.75 bar, and 70 mm, respectively. The nozzle travel path and speed were the same as all other experiment groups as shown in Figure 4. The same image analysis process was adopted, and corresponding material distribution has been compared with the predicted results of empirical model as shown in Figure 11. As can be seen in Figure 11, the proposed empirical model can well predict the material distribution of 3D spray cementitious material printing when the travel speed is 50, 100, and 250 mm/s. However, the predicted maximum filament widths are not accurate for cases when travel speed is 150 and 200 mm/s.1999Figure 11 Comparison between validation experimental and predicted material thickness distribution7. ConclusionIn this paper, the material distribution in 3D spray cementitious material printing was investigated.

ConclusionIn this paper, the material distribution in 3D spray cementitious material printing was investigated. With the adoption of robotic arm control, effects of four printing parameters (i.e. cementitious material flow rate, air flow rate, nozzle travel speed and standoff distance) to material distribution were studied. The experimental results showed that the increases of pumping speed or standoff distance increases the maximum filament width, the increase of nozzle travel speed decreases the maximum filament width, the change of air inject pressure has negligible effect on material distribution.

When the ratio between pumping speed and nozzle travel speed decrease to under some specific value, which requires further study, the sprayed filament density will decrease and thus, the cross section area will only decrease slightly.Based on the experimental material distribution results, an empirical model was proposed to describe the material distribution in 3D spray cementitious material printing. Different from the conventional concrete spray process, the influence of moving nozzle and2000extra brought by air flow with projected material have been taken into consideration. This empirical model predicts that the maximum filament width is linearly related to the product of standoff distance and the square root of the ratio of pumping speed and nozzle travel speed, then the material distribution is described by an empirical trapezoid shape model.

The verification experiments show that this empirical model can predict the maximum filament width and material distribution in a reasonable level. In the future, this empirical model will be improved and used in planning of 3D spray cementitious material printings.AcknowledgementThis research is supported by the National Research Foundation, Prime Minister’s Office, Singapore under its Medium-Sized Centre funding scheme, Singapore Centre for 3D Printing and Sembcorp Design & Construction Pte Ltd.Reference[1][2][3][4][5][6][7][8][9]C. K. Chua and K. F. Leong, 3D Printing and Additive Manufacturing: Principles and Applications: World Scientific Publishing Co Inc, 2017. B. Khoshnevis, "Automated construction by contour crafting—related robotics and information technologies," Automation in Construction, vol. 13, pp. 5-19, 2004. S. Lim, R. A. Buswell, T. T. Le, S. A. Austin, A. G. F. Gibb, and T.

Thorpe, "Mix design and fresh properties for high-performance printing concrete," Materials and Structures, vol. 45, pp. 1221-1232, 2012. B. Khoshnevis, X. YUAN, B. Zahiri, J. Zhang, and B. Xia, "Deformation Analysis of Sulfur Concrete Structures Made by Contour Crafting," in AIAA SPACE 2015 Conference and Exposition, Pasadena, USA, 2015, p. 4452. A. Perrot, D. Rangeard, and A. Pierre, "Structural built-up of cement-based materials used for 3D-printing extrusion techniques," Materials and Structures, vol. 49, pp. 1213- 1220, 2016. Y. Zhang, Y. Zhang, G. Liu, Y. Yang, M. Wu, and B. Pang, "Fresh properties of a novel 3D printing concrete ink," Construction and Building Materials, vol. 174, pp. 263-271, 2018.[10] Y. Weng, M. Li, M. J. Tan, and S. Qian, "Design 3D printing cementitious materials via Fuller Thompson theory and Marson-Percy model," Construction and Building Materials, vol. 163, pp. 600-610, 2018.[11] B. Panda, S. C. Paul, L. J. Hui, Y. W. D. Tay, and M. J.

163, pp. 600-610, 2018.[11] B. Panda, S. C. Paul, L. J. Hui, Y. W. D. Tay, and M. J. Tan, "Additive manufacturing of geopolymer for sustainable built environment," Journal of Cleaner Production, vol. 167, pp. 281-288, 2017.[12] D. G. Soltan and V. C. Li, "A self-reinforced cementitious composite for building-scale3D printing," Cement and Concrete Composites, vol. 90, pp. 1-13, 2018.2001[13] Y. Weng, M. Li, M. J. Tan, L. He, and S. Qian, "3D Printable High Performance Fiber Reinforced Cementitious Composites for Large-scale Printing," in Proceeding of the 3rd International Conference on Progress in Additive Manufacturing, Singapore, 2018, pp. 19-24.[14] W. Lao, M. Li, L. Masia, and M. J. Tan, "Approaching Rectangular Extrudate in 3D Printing for Building and Construction by Experimental Iteration of Nozzle Design," in Proceedings of Solid Freeform Symposium, Austin, TX, USA, 2017, pp. 2612-2623.[15] F. Bos, R. Wolfs, Z. Ahmed, and T.

2612-2623.[15] F. Bos, R. Wolfs, Z. Ahmed, and T. Salet, "Additive manufacturing of concrete in construction: potentials and challenges of 3D concrete printing," Virtual and Physical Prototyping, vol. 11, pp. 209-225, 2016.[16] B. Panda, S. C. Paul, N. A. N. Mohamed, Y. W. D. Tay, and M. J. Tan, "Measurement of tensile bond strength of 3D printed geopolymer mortar," Measurement, vol. 113, pp. 108-116, 2018.[17] ACI Committee 506, "Guide to shotcrete / reported by ACI Committee 506 (ACI 506R-05)," ed. Detroit, USA: American Concrete Institute, 2005.[18] B. Lu, M. Li, S. Qian, K. F. Leong, and M. J. Tan, "Develop Cementitious Materials Incorporating Fly Ash Cenosphere for Spray-based 3D Printing," in Proceedings of the 3rd International Conference on Progress in Additive Manufacturing, Singapore, 2018, pp. 38-43.[19] C. ASTM, "1437-01. Standard test method for flow of hydraulic cement mortar," ASTMInternational, Pennsylvania, 2001.[20] N. Ginouse and M.

38-43.[19] C. ASTM, "1437-01. Standard test method for flow of hydraulic cement mortar," ASTMInternational, Pennsylvania, 2001.[20] N. Ginouse and M. Jolin, "Mechanisms of placement in sprayed concrete," Tunnellingand Underground Space Technology, vol. 58, pp. 177-185, 2016.2002.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperThe Effect of Stress Intensity Factor on Fatigue Life of AM Parts Made from PolymerH. H. El Fazani1*, J. D.A. Coil2, R. R. Shah3, J. F. Laliberte4Department of Mechanical and Aerospace Engineering, Carleton University,1125 Colonel By Drive, Ottawa, Ontario, K1S-5B6, Canada.Corresponding Author (hayatelfazani@cmail.carleton.ca)AbstractThe fatigue characteristics of additively manufactured specimens was investigated. A commercial acrylonitrile-butadiene-styrene (ABS) polymer (P430) was selected to manufacture AM fatigue coupons due to its low cost and wide applications. A total of 30 fatigue coupons were built on flat and on edge using a Stratasys SST 1200es fused deposition machine. The AM samples were manufactured at different build orientations of 0°, 22.5°, 45°, 67.5°, and 90°.

The AM samples were manufactured at different build orientations of 0°, 22.5°, 45°, 67.5°, and 90°. The objective of this experiment is to investigate the influence of stress intensity factor on fatigue life of AM polymer parts. The specimens were tested under low sinusoidal tension-tension fatigue loading. The fatigue crack behaviour was monitored using a travelling microscope method. The stress intensity factor was investigated as the cycles accumulated. The effect of stress intensity factor on fatigue life was examined. The investigation of fatigue crack growth as a function of the number of fatigue cycles was discussed. It was found the fatigue coupons manufactured at 0° build orientation have a better fatigue life compared with the other build orientations. More investigation is required for other AM parameters such as layer thickness, infill density, manufacturing process to help better understand the fatigue performance of AM part made by polymer material.

The Scanning Electron Microscope (SEM) technology was used to characterize the fractured surfaces and void distribution of AM fatigue parts. The void percentage was estimated. The results showed that AM fatigue parts built on flat has a higher void percentage compared to the AM parts manufactured on edge.KeywordsAdditive Manufacturing, Fused deposition modelling, Fatigue Crack, Fatigue Coupons, Stress Intensity Factor.624IntroductionAdditive manufacturing technology has brought great advantages. However, there is a current need to further study the material behavior of AM parts. The lack of research publication on fatigue analysis for AM polymer parts manufactured by FDM technique was discussed in [1] [2]. For this reason, the fatigue crack propagation behavior for polymers continues to be an active area of study.[3]. Lee et al. [4] explored the influence of fatigue on AM part made from ABS material.

For this reason, the fatigue crack propagation behavior for polymers continues to be an active area of study.[3]. Lee et al. [4] explored the influence of fatigue on AM part made from ABS material. The study was based on comparing the fatigue life of ABS polymer made by FDM with a bulk precursor material. The results demonstrated that AM polymers have a significantly lower fatigue life. Likewise, it is crucial to consider long term fatigue loading for aircraft applications; ABS 3D printed polymers have very poor fatigue characteristics, which limits their feasibility for application in the aircraft structure, concluded by Schiller [5].It is necessary to understand the damage propagation mechanics underlying the failure process to ensure the final AM product will not suffer early failure under cyclic loading. Fatigue crack growth testing was conducted to examine both initiation and propagation mechanisms, with a focus on early crack initiation in AM polymer materials.

Fatigue crack growth testing was conducted to examine both initiation and propagation mechanisms, with a focus on early crack initiation in AM polymer materials. Several fatigue testing methodologies are available and described in the literature [1][4] [6]. The use of a specific method depends on the types of the fluctuating stresses and strains, as well as on the purpose of test. Fatigue tests can be based on reversed stress cycles, repeated stress cycles, or random or irregular stress cycles. These methods are also known as tension – compression fatigue loading, compression – compression fatigue loading, and tension – tension, respectively. Testing for each can be conducted at high cyclic fatigue loading, or low cyclic fatigue loading. The method which involves the opening in tension fatigue mode I was chosen in this experiment. The fatigue test approach that was utilized in this research was based on a tension – tension, low cyclic fatigue loading as shown in Figure 1.

The fatigue test approach that was utilized in this research was based on a tension – tension, low cyclic fatigue loading as shown in Figure 1. The traveling microscope approach was selected for crack detection and monitoring, and crack measurement. The method was chosen because the traveling microscope gives accurate crack measurements both at initiation and during propagation. In this research, the application of damage tolerance and fatigue evaluation and fatigue characteristics of ABS material made by FDM were investigated. This includes investigation of the influence of the stress intensity factor on fatigue life. The fatigue crack growth in AM polymer specimens were monitored and the crack lengths were measured by using a travelling microscope.Figure 1: Constant amplitude fatigue.625MethodologyThe fatigue life of AM polymer parts was determined by measuring the crack length on AM specimens caused by low cyclic, tension-tension fatigue loading.

The fatigue failure of AM parts starts with a crack formation at a point of high stress concentration. This is followed by crack propagation, which ultimately leads to total failure of the component. An optical microscope was used to monitor the crack lengths at initiation and propagation. Table 1 shows the test matrix for fatigue analysis using the travelling microscope. The test matrix was made to investigate the effect of building orientations 0°, 22.5°, 45°, 67.5°, and 90° on the fatigue properties of AM polymers. The parts were manufactured with 0.25 mm layer thickness.

The parts were manufactured with 0.25 mm layer thickness. The minimum number of proposed fatigue tests for travelling microscope method is 30 tests.Table 1: Material fatigue test matrix.AM process and orientationsMinimum number of acceptable fatigue testTotalStratasys SST 1200esFlat0° 422.5° 45° 67.5° 90° 734321Layer Thickness 0.25 mmOn Edge3339Total30Fatigue Test Setup and ProcedureThe basic fatigue test consists of a test coupon that is subjected to a cyclic load. The frequency and stress amplitude applied during fatigue testing will cause the polymer to exhibit either thermal softening or cyclic crack growth [7]. The AM fatigue specimen was designed as a large radii dogbone coupon, reducing the possibility of premature specimen failure during cyclic loading while still maintaining a small region of possible crack formation. The MTS machine used for the fatigue testing is shown in Figure 2a.

The MTS machine used for the fatigue testing is shown in Figure 2a. In this research, the specimens were tested under sinusoidal tension-tension fatigue loading at a low frequency of 3 Hz. The maximum applied load was + 0.9 kN and the minimum applied load was zero. The selection of this type of loading condition was based on previous literature [4][8]. Testing the specimen at low frequency eliminates any heating effects due to the strains. As well, using a minimum applied load of zero avoids any bending that might occur from compression. In addition, if a compression load is applied, it becomes very difficult to accurately identify the crack length as the crack will be closed during compression. The undesired compression force may also lead to losing the features of the fracture surface, which are very important to the analysis. A travelling microscope was utilized to observe and measure the cracks with an accuracy of +/- 0.01 mm.

A travelling microscope was utilized to observe and measure the cracks with an accuracy of +/- 0.01 mm. A point light source was used to illuminate the specimen surface. The specific type of light was chosen to clearly identify cracks throughout the test. The specimen was gripped axially in the loading direction as shown ion Figure 2b.626Load frameLightAlignementTravellingi(a) (b)Figure 2: Fatigue test equipment, a. MTS machine, b. Fatigue test specimen aligned fences.The fatigue test was conducted at room-temperature (approximately 21°C). The specimen surface was observed during loading and unloading. A small crack which started to form and initiate on the gauge length of the tested area was tracked. There are different techniques of crack detection such as the compliance method applied in [9].

A small crack which started to form and initiate on the gauge length of the tested area was tracked. There are different techniques of crack detection such as the compliance method applied in [9]. The chosen method of crack monitoring was to use travelling microscope, where the crack tip was monitored with a crosshair and with light point assistant as shown in Figure 3. The microscope travels parallel to the crack using a hand cranked vernier mechanism. The travelling microscope has a digital scale readout, allowing a crack measurement within an accuracy of 0.01 mm [10]. The crack propagation was monitored until the crack reached critical dimensions or until a complete specimen failure occurred [11]. The visual microscope approach is considered a reliable method because it gives crack measurements accurate enough to be used to validate other techniques [9]. The first specimen was cycled for 300 cycles.

The first specimen was cycled for 300 cycles. Once this point was reached, the gauge area was carefully observed by using the travelling microscope to identify the cracks. All cracks and their respective lengths were measured and recorded. Images in Figure 3 shows all the cracks formed both on and inside the surface. After crack initiation, continual loading of the stress concentration with additional cycles led to a gradual crack propagation. The crack accelerates over more cycles, which sometimes culminates a sudden failure. These steps were repeated until the specimen failed. The number of cycles at the failure was recorded and all data was saved for analysis. The number of cycles associated with the crack measurements was recorded with the applied loads during the measuring. The speed of an individual crack propagation can be influenced by other cracks surrounding a single crack.

The speed of an individual crack propagation can be influenced by other cracks surrounding a single crack. It is normal for multiple cracks under cyclic stress fluctuations to grow and come together to form a single long crack. The crack then becomes unstable, resulting in a sudden quick failure which separates the specimen into two parts. SEM analysis was conducted for the fracture surfaces of AM fatigue coupons to investigate the void distribution and to calculate the void volume.627Figure3: Fatigue testing procedure: a. Specimen mounted at 25 kN MTS load frame, b. Under fatigue testing, c. Light source, d. Travel microscope, e. Multiple cracks, f. Single crack length measurement, g. Specimen at failure.Fatigue Crack Growth Rates as a Function of Stress-Intensity Factor Range (ΔK)Fatigue crack growth rates was uniquely correlated to stress intensity factor range (ΔK) by Furmanski et al. [3].

[3]. In literature there are several approaches to determine the cyclic crack growth rate, da/dN , where a is the typical crack length and N is the number of fatigue cycles [9][10] [12] [13]. For example, it was suggested by Bannantine et al. [14] and Blattmeier et al.

For example, it was suggested by Bannantine et al. [14] and Blattmeier et al. [15] that the fatigue crack propagation rates under cyclic loading can be empirically represented by Equation 1, which also known as Paris power law.𝑚𝑚 (1) represents the stress intensity factor range in MPa m1/2, C and m are published material constants [11][16] found empirically, showing correlation between the crack growth rate with the amplitude cyclic stress intensity.𝑑𝑑𝑑𝑑 𝑑𝑑𝑑𝑑= 𝐶𝐶(∆𝐾𝐾)Where∆𝐾𝐾628By using the crack measurements recorded from the travelling microscope and the number can beof cycles N displayed in the MTS software, the ranges of stress intensity factor calculated by:∆𝐾𝐾(2)The maximum load level of 0.9 kN was specified and the minimum load selected to be ∆𝐾𝐾 = 𝐾𝐾𝑚𝑚𝑚𝑚𝑚𝑚 − 𝐾𝐾𝑚𝑚𝑚𝑚𝑚𝑚 = 𝑓𝑓(𝑔𝑔)∆𝜎𝜎√𝜋𝜋𝑑𝑑zero.

It is important to highlight that Paris law is observably identified in this curve. The specimen was cycled at low cyclic loading and the stress intensity factors were estimated as the cycles were accumulating. It was observed that over time the crack accelerates, and the crack speed increased after reaching a stress intensity value of 0.65 (MPa m1/2). To get a better understanding on how the build orientations for AM part may affect the crack growth rates. The results of fatigue crack growth rate as a function of stress intensity factor for a group of specimens manufactured with 0°, 22.5°, 45°, 67.5° and 90° build orientations are illustrated in Figure 6. This graph provides a summary of the crack growth rates (da/dN) for 5 AM specimens. A rapid growth in crack length was observed for AM specimen manufactured with6290° orientation. The crack growth rate reached 0.0015 (m/cycle) at a stress intensity factor value of 0.62 (MPa m1/2).

The crack growth rate reached 0.0015 (m/cycle) at a stress intensity factor value of 0.62 (MPa m1/2). It was expected for part built with 90° to have the same behavior but this part exhibits a slower propagation in which that crack growth rates was 0.0006 (m/cycle) at stress intensity factor of nearly 0.77 (MPa m1/2). The fatigue crack growth rate for specimens built on 67.5° exhibited the lowest amount in comparison with other specimens. There is clear overlap between the final stage of crack initiation for the part built at 90°.

There is clear overlap between the final stage of crack initiation for the part built at 90°. The specimen with a build orientation of 45° demonstrated a gradual increase in crack length.0.00070.0006) e l c y c / m(0.00050.00040.0003N d / a d0.00020.0001000.20.4 0.6 ΔK (MPa m1/2)0.81Figure 5: Fatigue crack growth rate for a single crack as a function of stress-intensity factor range (ΔK) for AM specimen manufactured on flat and with 90° build orientations.0.0016) e l c y c / m0.00140.00120.0010.00080 deg 22.5 deg 45 deg 67.5 deg 90 deg(N d / a d0.00060.00040.0002000.20.40.60.81ΔK (MPa m1/2)Figure 6: Fatigue crack growth rate for a single crack as a function of stress-intensity factor range (ΔK) for specimens manufactured on flat and with 0°, 22.5°, 45°, 67.5° and 90° build orientations.The relation between the fatigue crack growth rate (da/dN) and the crack length (a) for a single crack is shown in Figure 7.

The presented results are for AM parts made with 0°, 22.5°, 45°, 67.5° and 90° build orientations. It was observed that the AM part manufactured with 0° can survive larger cracks before failure compared to the other orientations.6300.0030.0025) e l c y c / m(N d / a d0.0020.00150.0010° 22.5° 45° 67.5° 90°0.0005000.00050.0010.00150.0020.0025a (m)Figure 7: Fatigue crack growth rate for a single crack as a function of crack length for specimens manufactured on flat and with 0°, 22.5°, 45°, 67.5° and 90° build orientations.Crack Growth Rates Analysis for Specimens Manufactured with on edgeThe results of fatigue crack growth rate versus the stress intensity factor for AM part manufactured with 90° build orientations are displayed in Figure 8. The crack growth rate of 3.42 *10-5 (m/cycle) was recorded at stress intensity factor of 0.617 (MPa m1/2).

The crack growth rate of 3.42 *10-5 (m/cycle) was recorded at stress intensity factor of 0.617 (MPa m1/2). The crack grew slowly until it reached a rate of 1.1*10-4 (m/cycle) at an approximate stress intensity factor of 0.94 (MPa m1/2). Then the crack speed increased which led to rapid crack propagation. The final stress intensity factor collected is 2.11 (MPa m1/2) and the crack growth rate was measured to be 1.83*10- 4 (m/cycle).) e l c y c / m(N d / a d0.0002 0.00018 0.00016 0.00014 0.00012 0.0001 0.00008 0.00006 0.00004 0.00002 000.511.522.5ΔK (MPa m1/2)Figure 8: Fatigue crack growth rate for a single crack as a function of stress-intensity factor range (ΔK) for AM specimen manufactured on edge and with 90° build orientations.631Figure 9 shows the results of fatigue crack growth rate versus the stress intensity factor for AM fatigue specimens manufactured on edge and with build angles of 0°,45° and 90°.

The AM coupon built with 0° orientations exhibited a higher crack growth in comparison with the 45° and 90° specimens. The highest crack propagation was recorded at 2.13 (MPa m1/2) with a value of da/dN of 3*10-4 (m/cycle). It was also observed that there is an overlap between the AM samples printed at 45° and 90° build angles. This overlap was found within a range of stress intensity factors from 0.75 to 2.11 (MPa m1/2). Since the crack growth rate da/dN for AM fatigue coupon built on 45° was found to be as low as 0.001 at 3.5 (MPa m1/2).

Since the crack growth rate da/dN for AM fatigue coupon built on 45° was found to be as low as 0.001 at 3.5 (MPa m1/2). Therefore, it can be argued that the AM part oriented at 45° has a better fatigue life compared to the coupons oriented at 0° and 90° build orientation.0.010.009) e l c y c / m(0.0080.0070.0060.005N d / a d0.0040.0030° - 1 45° - 2 90° - 30.0020.001000.511.522.533.5ΔK (MPa m1/2)Figure 9: Fatigue crack growth rate for a single crack as a function of stress-intensity factor range (ΔK) for specimens manufactured on edge and with 0°, 45° and 90° build orientations.Scanning Electron Microscope (SEM) ResultsThe SEM examination of the fractured surface of AM polymer parts built on flat at 45° build orientations is shown in Figure 10. The surface labeled 1 was close to the building bed and it was built first, surface 2 was built at the end.

The surface labeled 1 was close to the building bed and it was built first, surface 2 was built at the end. It is contended that the indicated fracture path followed the weakest portions of the sample. The voids and the cavities formation and interconnection over cyclic loading increased the overall crack length and led to sudden failure. Figure 11 displays the SEM analysis for AM part made on edge side at 90° build orientation using Stratasys SST 1200es. The surface labeled 1 was closest to the heated building bed and was deposited first, surface 2 was deposited at the end. The microscopy results show that the filaments in the middle of the cross-section area had better fusion which led to fewer airgaps. It is likely that the specimens experienced a brittle fracture. The fatigue crack features on the surfaces were also highlighted.

It is likely that the specimens experienced a brittle fracture. The fatigue crack features on the surfaces were also highlighted. A voids distributionAM parts manufactured on flat found to have a higher void percentage of 7.29% compared to the fatigue coupons built on edge with voids percentage 2.26%. This finding support the fatigue crack growth results discussed in the previous section.

This finding support the fatigue crack growth results discussed in the previous section. In other words, the fatigue crack growth632rate was higher for the parts manufactured on flat versus on edge which indicates that the void formation has a critical impact on the fatigue life.VoidsFigure 10: SEM image for the fractured surface of AM fatigue coupon made on the flat side at 45° build orientations using Stratasys SST 1200es.VoidsFigure 11: SEM microscopy for the fractured surface of AM fatigue coupon manufactured on edge at 90° build orientation using Stratasys SST 1200es.633ConclusionFatigue crack investigation was conducted for AM parts made from ABS polymer material. The influence of the stress intensity factor on the fatigue life of AM fatigue coupons was examined. The fatigue crack measurements results for AM polymer specimens showed that the building orientation and the existence of voids within the specimens played a significant role in the fatigue properties.

The coupons built at 0° orientation can accumulate more damage and thus have a longer fatigue life as compared to the other four orientations 22.5°, 45°, 67.5° and 90°. SEM analysis was performed for the fractured surface of AM polymer parts. The voids percentage of AM fatigue parts were calculated. The AM parts built of flat found to have a higher voids percentage compared to the AM parts manufactured on edge.In future work, it is recommended that expand the optical analysis to include Micro Computed Tomography (Micro-CT). Deep investigation is required for the internal structure of AM fatigue parts manufactured at 0° build orientation. This will provide a better understanding the reasons for accumulating longer crack for this specific samples.AcknowledgementsThe authors would like to acknowledge the technical support and training provided by the supervisor of the Mechanical and Aerospace Engineering Laboratories at Carleton University Mr.

fect of Surface Preparation Methods on Mechanical Properties of 3D Structures Fabricated by Stereolithography and 3D Printing for Electroless Ni PlatingShine Joseph, Stella Quiñones, Frank Medina, Ryan Wicker W.M. Keck Center for 3D Innovation University of Texas at El Paso El Paso, Texas 79968Abstract Stereolithography (SL) and 3D Printing (3DP) are useful technologies for three-dimensional prototyping applications, providing highly accurate and detailed part geometries with high quality surface finishes. It is desired to improve the materials performance of the existing photocurable SL and 3DP resins for rapid tooling and other functional applications by applying a nickel (Ni) coating. In this work, surface preparation methods for electroless plating of commercial photopolymer resins such as NanoFormTM15120 (NanoForm) and Objet FullCure®840 (Veroblue) were explored in order to enhance the structural integrity of RP components.

This study examined different surface preparation methods (chemical etching) and their effect on the surface morphology and mechanical strength of the polymers. It was observed that surface preparation of the resins significantly affected the mechanical properties and Ni plating of the substrate polymers. This is a critical step, since the Ni film takes on the surface structure of the substrate.Keywords: Stereolithography, 3D Printing, Electroless Ni Plating, Etching, NanoForm, VeroBlue.1. Introduction Rapid prototyping (RP) is a processing technology where a three dimensional model is manufactured as a result of a computer aided design (CAD) file that is sliced and solidified one layer at a time. There are several RP technologies such as stereolithography (SL), 3D Printing (3DP), fused deposition modeling (FDM) and selective laser sintering (SLS). Advances in RP technology have contributed greatly to the development phase in the design process [1].

Advances in RP technology have contributed greatly to the development phase in the design process [1]. While the contribution to the manufacturing arena is significant, the process itself is not considered a true manufacturing process, since there are limitations associated with dimensional accuracy and the mechanical properties of RP components. Metallization of RP polymers is a way to enhance the mechanical properties for rapid tooling, and possibly other end-use applications. A study by Saleh et al. (2004) concluded that the ultimate tensile strength and Young’s Modulus was improved for SL and SLS parts plated with copper and nickel, with less of an effect on % elongation. The plating also helped to improve the surface roughness of SLS components [2].

The plating also helped to improve the surface roughness of SLS components [2]. The plating of RP parts could prove to be a useful tool for SL materials where the long term behavior of properties can be predicted.An additional application is the selective deposition of metallic patterns or conductive lines that can be used for electronic interconnects for embedded circuits, where the electronic components are embedded in a polymer matrix and conductive lines are printed by various methods. Examples of this technology include direct write (DW) and selective electroless copper (Cu) and nickel (Ni) deposition [3-7]. RP processes are especially useful when manufacturing 3D circuits of arbitrary shape and size.In this paper, the electroless Ni deposition process was examined, specifically the surface enhancing step. Electroless Ni plating involves several processing steps, including etching of the polymer substrate in order to create a strong bond at the metal/polymer interface.

Electroless Ni plating involves several processing steps, including etching of the polymer substrate in order to create a strong bond at the metal/polymer interface. Since etching requires property changing chemicals, such as strong acids, the effect of this step was examined for two RP resins. The ultimate tensile strength and % elongation of NanoFormTM15120 (NanoForm), an SL resin, and Objet FullCure®840 (Veroblue), a 3DP resin, were examined using two etchants at various concentrations. The mechanical properties were correlated with optical microscopy before and after etching in order to determine the overall interaction between the substrate and the etchant.2. BackgroundElectroless plating is a catalytically driven electrochemical reduction process where no electricity is required. In electroless Ni plating, Ni ions (Ni2+) in solution are reduced by a reducing agent on the surface of a catalytically activated substrate.

In electroless Ni plating, Ni ions (Ni2+) in solution are reduced by a reducing agent on the surface of a catalytically activated substrate. It has been observed that the surface preparation of the substrate is the crucial phase in all electroless plating processes.The electroless Ni plating process includes five major steps which are (1) etching, (2) neutralization, (3) sensitization, (4) surface activation and (5) electroless Ni plating. The substrate surface needs to be roughened (chemical or mechanical etch) appropriately to facilitate mechanical anchoring of the metal on the substrate. This step improves the bonding of the metal film to the polymer substrate. In a hypophosphite electroless nickel bath, hypophosphite ions reduce Ni ions in the presence of a catalyst on the substrate surface. Low cost is one of the advantages of hypophosphite over other reducing agents. Several mechanisms have been proposed for Ni reduction by hypophosphite.

Low cost is one of the advantages of hypophosphite over other reducing agents. Several mechanisms have been proposed for Ni reduction by hypophosphite. The most widely accepted mechanism of Ni reduction is through the oxidation of sodium . H2O (Rxn 1). The further oxidation of hypophosphite yields hypophosphite, NaH2PO2 electrons for Ni and hydrogen reduction at the catalytically activated surface (Rxns. 2, 3 and 4). The overall reaction for the process is shown in Rxn 5 [8-10].Dissociation of Hypophosphite: Na+ + H2PO2 NaH2PO2−(Rxn. 1)Oxidation of Hypophosphite: − + H2O H2PO3 H2PO2− + 2H+ + 2e−(Rxn. 2)Nickel Reduction: Ni++ + 2e− Ni(Rxn. 3)Hydrogen Reduction: 2H+ + 2e− H2(Rxn. 4)Overall Reaction: Ni++ + H2PO2− + H2ONi + H2PO3− +2H+(Rxn. 5)3. Experimental Procedure NanoForm and VeroBlue resins were chosen for this study since electroless Ni deposition occurs readily on these surfaces, especially VeroBlue.

5)3. Experimental Procedure NanoForm and VeroBlue resins were chosen for this study since electroless Ni deposition occurs readily on these surfaces, especially VeroBlue. NanoForm is an SL nanoparticle filled material with high stiffness SL properties, and VeroBlue is a 3D Printing material known for its impact resistance. There is a particular interest in NanoForm, since this material is being used as a substrate for the deposition of narrow conductive channels of electroless Ni. These channels can be used to connect electronic components for 3D circuits and for antenna applications.The electroless Ni process involves several chemical processes, and this study focuses on the etching of the photocurable resin. Etching of the polymer surface is necessary in order to create an even microporous surface that promotes bonding with the metal catalyst and film. ASTM D638 Type I NanoForm and VeroBlue tensile specimens were built using SL and 3DP, respectively.

ASTM D638 Type I NanoForm and VeroBlue tensile specimens were built using SL and 3DP, respectively. The tensile samples were etched and tested in order to determine the effect of the etching process on mechanical properties. Two etchants, chromic acid and potassium permanganate, were used since they have different surface altering characteristics. Table 1 includes the design of experiments with 4 etchant compositions and 3 etch times for each composition. The design of experiments is based on preliminary work to determine appropriate composition and etch time values. The surfaces of these samples were examined optically before and after etching, in order to correlate degree of etching with the presence and/or quality of the Ni film. This is necessary in order to optimize the plating process for both planar and selective film deposition.Table 1.

This is necessary in order to optimize the plating process for both planar and selective film deposition.Table 1. Design of Experiments for NanoForm and VeroBlue Resins with Etchant Composition and Etch TimeEtchantConcentration (g/L)Etching Time (minutes)40Potassium Permanganate (KMnO4)50605101570300Chromic Acid (CrO3)32535025103753.1 Nanoform Sample Build Procedure NanoForm tensile samples were fabricated by stereolithography using a Viper Si2 machine in a flat orientation. A layer thickness of 0.004″ (0.1 mm) was used with a critical exposure (Ec) of 16.3 mJ/cm2 and a cure depth (Dp) of 0.132 mm. The samples were cleaned in isopropyl alcohol after manually removing supports and dried in compressed air. Post curing was performed in a UV oven for 30 minutes on each side. Five ASTM D638 Type 1 tensile samples were fabricated for each experimental condition (Fig. 1).

Post curing was performed in a UV oven for 30 minutes on each side. Five ASTM D638 Type 1 tensile samples were fabricated for each experimental condition (Fig. 1). Table 2 includes the dimensions and tolerance values for this standard for each critical dimension. All samples were carefully measured using a Mitutoyo® 573 series caliper in order to ensure ASTM standard dimensions.Figure 1. ASTM D638 Type I Sample Illustrating Critical Dimensions [11]Table 2. ASTM D-638 Type I tensile test specimen dimensions and tolerances [11]Dimensions (in)W, (Wc) - Width of narrow section L - Length of narrow section Wo - Width overall Lo - Length overall G - Gage length D - Distance between grips R - Radius of fillet T - ThicknessType I 0.5 2.25 0.75 6.5 2 4.5 3 0.13Tolerances ± 0.02 ± 0.02 + 0.25 no max ± 0.01 ± 0.2 ± 0.04 ± 0.023.2 VeroBlue Sample Build ProcedureVeroBlue samples were built by 3D printing using an Eden 333 machine.

The samples were cleaned in water after removing supports and dried in compressed air. During the build process, the material was cured layer by layer with UV lamps, which eliminated the need to further cure the material. The layers were jetted out and cured with layer thickness values of 0.6 mils (16 µm).3.3 Electroless Ni Plating ProcedureThe electroless plating procedure used in this study can be divided into two main steps, surface preparation and plating. The surface preparation steps include etching, neutralization, sensitization and surface activation.EtchingNeutralizationSensitizationCatalytic activationElectroless nickel platingThe two etchants chosen for this study were chromic acid (CrO3) and potassium permanganate (KMnO4). The chromic acid etchant was a combination of 21 mL/L of concentrated sulfuric acid and chromic acid concentrations of 300, 325, 350 and 375 g/L in deionized water (DI) at a temperature of 60–65oC.

The chromic acid etchant was a combination of 21 mL/L of concentrated sulfuric acid and chromic acid concentrations of 300, 325, 350 and 375 g/L in deionized water (DI) at a temperature of 60–65oC. The etch times used for the chromic acid etchant were 2, 5 and 10 minutes (Table 1). The potassium permanganate etchant contained 60g/L sodium hydroxide (NaOH) and a concentration series of 40, 50, 60 and 70g/L of potassium permanganate dissolved in DI water at 80–85oC (Table 1). The samples were rinsed two times in DI water after each step in the electroless Ni plating process.The samples were etched and neutralized in order to remove the excess etchant from the substrate surface. The chromic acid etched samples were rinsed in DI water and neutralized in 2% NH4OH for 30 seconds. The samples etched in potassium permanganate were immersed in oxidant sulfuric acid solution (2% vol. H2SO4 and 2% vol. H2O2) for 5 minutes.

The samples etched in potassium permanganate were immersed in oxidant sulfuric acid solution (2% vol. H2SO4 and 2% vol. H2O2) for 5 minutes. The samples were then sensitized (pre-activation) with a solution of 80 g/L of stannous chloride (SnCl2) and 45 mL/L of concentrated hydrochloric acid (HCl) dissolved in DI water at room temperature. The sensitized substrate surface was activated with palladium catalyst at room temperature. The catalyst solution contained 1 g/L of palladium chloride (PdCl2), 45 mL/L of concentrated hydrochloric acid and deionized water.The electroless Ni plating bath consisted of 30 g/L of nickel chloride (NiCl2. 6H2O), 10 g/L sodium glycolate (HOCH2CO2Na) and 10 g/L sodium hypophosphite (NaH2PO2. H2O). The catalytically activated samples were transferred carefully to the electroless Ni plating bath maintained at a temperature of 80–85oC with oxygen circulation. The constant temperature of the plating bath was maintained in a circulating water bath.

The constant temperature of the plating bath was maintained in a circulating water bath. The oxygen circulation in the bath was used to avoid the possibility of spontaneous decomposition of the hypophosphite bath at this temperature. The samples were plated for approximately 5 minutes and removed from the plating bath and cooled to room temperature and cleaned in DI water.3.4 Mechanical Testing of NanoForm and VeroBlueMechanical testing was carried out using an Instron 5866 tensile testing machine. Tensile testing of the NanoForm and VeroBlue samples (ASTM D638, Type 1) were conducted before and after etching. A set of 5 tensile specimens were tested for NanoForm and VeroBlue for each of the design parameters listed in Table 1. The tensile samples were neutralized and rinsed prior to mechanical testing. A comparison was made between the two materials based on etchant type, concentration and etch time.4.

The tensile samples were neutralized and rinsed prior to mechanical testing. A comparison was made between the two materials based on etchant type, concentration and etch time.4. Results and Discussion4.1 Etching and Plating of NanoForm and VeroBlueFigure 2(a) and (b) are optical micrographs illustrating the surface view of NanoForm and VeroBlue prior to etching. There are distinct differences between the two materials, NanoForm is relatively smooth, while VeroBlue includes obvious layers as a result of the 3D print build process. The exposure of these resins to the chromic acid and potassium permanganate etchants altered the surfaces in a distinct manner that can be attributed to the type of etchant. Plating was successful for most of the samples etched with chromic acid, while the success with the potassium permanganate etchant was sporadic as indicated in Tables 3 and 4.

Plating was successful for most of the samples etched with chromic acid, while the success with the potassium permanganate etchant was sporadic as indicated in Tables 3 and 4. In all cases where plating occurred, the electroless Ni film took on the surface features of the substrate.45µm(a)(b)Figure 2. Optical view of (a) NanoForm and (b) VeroBlue Samples Prior to Etching4.1.1 Chromic Acid Etchant (NanoForm and VeroBlue Resins)The NanoForm and VeroBlue samples were etched with chromic acid at concentrations of 300, 325, 350 and 375 g/L for 2, 5 and 10 minutes at each concentration. For the NanoForm samples, the chromic acid etchant resulted in a fine dispersion of uniform etch pits for most etchant concentrations for etch times less than 10 minutes. For all values of etchant concentrations, surface cracking was observed at etch times of 10 minutes, with some cracking visible at 5 minutes for the highestTable 3.

Characterization of Ni Film on NanoForm and VeroBlue Samples Etched with Potassium PermanganateEtching TimeEtchantConcen- tration5 minutesVeroBlue 10 minute s15 minutes5 minutesNanoForm10 minutes15 minutes40 g/LBrightDark – unevenDark – unevenIncom- pleteIncom- pleteUnevenPotassium Permang- anate (KMnO4)50 g/LBright - unevenDark – unevenDark – unevenIncom- pleteIncom- pleteBright – uneven60 g/LBrightDark – unevenDark – unevenIncom- pleteIncom- pleteIncom- plete70 g/LBrightDark – unevenDark – unevenIncom- pleteIncom- pleteIncom- pleteconcentration (375 g/L). In comparison, the Ni coatings on NanoForm for t = 10 minutes were inferior in comparison to lower etch times (Table 3).A very different effect was observed for the VeroBlue samples. These samples have a very rough surface morphology due to the 3D build process, where layers are evident.

These samples have a very rough surface morphology due to the 3D build process, where layers are evident. A reduction in the degree of visibility of the layers was observed with increase in etchant concentration. No selective attack on the surface was observed.Figures 3 and 4 include optical micrographs of NanoForm and VeroBlue, respectively, illustrating the plating morphology dependence on the chromic acid etchant surface modification step for concentrations of 350g/L (5 min etch) and 375 g/L (10 min etch). Cracking in the etched sample (Fig. 3b) is evident after plating as indicated by the arrows. As a result of chromic acid etching, high quality Ni films were observed on both materials, with slightly brighter films on the NanoForm surfaces (Table 3).EtchedPlated45µm(a)(b)Figure 3. Optical Micrographs of NanoForm Etched and Plated with (a) 350 g/L chromic acid for 5 minutes and (b) 375 g/L chromic acid for 10 minutesEtchedPlated45µm(a)(b)Figure 4.

Optical Micrographs of NanoForm Etched and Plated with (a) 350 g/L chromic acid for 5 minutes and (b) 375 g/L chromic acid for 10 minutesEtchedPlated45µm(a)(b)Figure 4. Optical Micrographs of VeroBlue Etched and Plated with (a) 350 g/L chromic acid for 5 minutes and (b) 375 g/L chromic acid for 10 minutes4.1.2 Potassium Permanganate Etchant (NanoForm and VeroBlue Resins) An evaluation of the second etchant consisted of potassium permanganate concentrations of 40, 50, 60 and 70 g/L and etch times of 5, 10 and 15 minutes. The potassium permanganate etchant resulted in a more severe attach on both the NanoForm and VeroBlue samples. Blistering was observed on most of the NanoForm samples, and became more pronounced with increase in concentration (Fig. 5). Very few of the NanoForm samples etched with permanganate were plated completely (Table 4).For VeroBlue, exposure to the permanganate etchant at most concentrations removed all evidence of the layering.

All of the VeroBlue samples were plated, and evidence of over-etching was visible on the Ni film (Fig. 6).Figure 5 includes optical micrographs for two NanoForm samples etched with 50 g/L (15 min) and 70 g/L (10 min) potassium permanganate and plated. The strong effect of etching is evident in the Ni film. Figure 6 shows a similar effect for VeroBlue at etchant concentrations of 40 g/L (5 min) and 60 g/L (10 min). As is evident in both Figs. 5 andEtchedPlated45µm(a)(b)Figure 5. Optical Micrographs of NanoForm Etched and Plated with Permanganate at (a) 50 g/L for 15 min and (b) 70 g/L for 10 min6, the Ni film structure mirrors the substrate features arising due to etching. Very poor films were observed by using the potassium permanganate etchant on both resins.5.2 Tensile Properties of Etched SamplesThe effect of etchant type, concentration and time was examined for ASTM D-638 Type I tensile samples of NanoForm and VeroBlue. The results are shown in Figs. 7-10.

The results are shown in Figs. 7-10. A comparison based on etchant type was made for each material for 4 values of concentration, and 3 etch times at each concentration. The notation C=0/t=0 indicates the resin material after RP build and post curing, without etchant treatment and is referred to as the base material.5.2.1 Chromic Acid Etchant (NanoForm and VeroBlue Resins)The ultimate tensile strength for NanoForm was relatively uniform for all concentration values, with an average 6% loss in strength due to chromic acid etching (Fig. 7). A slight increase in % elongation was also observed, independent of etchant concentration (Fig. 8). The average % elongation was 2.68% compared to 2.36% for theEtchedPlated45µm(a)(b)Figure 6. Optical Micrographs of VeroBlue Etched and Plated with Permanganate at (a) 40 g/L for 5 min and (b) 60 g/L for 10 minNanoForm base material.

Optical Micrographs of VeroBlue Etched and Plated with Permanganate at (a) 40 g/L for 5 min and (b) 60 g/L for 10 minNanoForm base material. As mentioned above, chromic acid resulted in a uniform attack on NanoForm at all concentrations, with some cracking at the highest etch times (10 minutes).The chromic acid had a more pronounced effect with respect to the VeroBlue base material, with an average loss in strength of 11%. As with NanoForm, the VeroBlue UTS was independent of chromic acid concentration, and the values ranged from 35-39 MPa compared to 42 MPa for the base material (Fig. 9). The VeroBlue samples became less ductile with etching, with an average % elongation of 8%, compared to 17% for the base material.

9). The VeroBlue samples became less ductile with etching, with an average % elongation of 8%, compared to 17% for the base material. A wide range of % elongation values were observed for this material, between 4.91 – 13%.5.2.2 Potassium Permanganate Etchant (NanoForm and VeroBlue Resins) A systematic decrease in UTS was observed for both NanoForm and VeroBlue etched with permanganate with increase in etchant concentration and etch time. For NanoForm,C = 0/t= 080.00C =1/t = 175.00C = 1/t =270.00C = 1/t= 3a P M,S T U65.0060.0055.00C = 2/t= 1C = 2/t= 2C = 2/t= 3C = 3/t= 150.00C = 3/t= 245.00C = 3/t= 340.00C = 4/t= 1KMnO4Chromic AcidC = 4/t= 2C = 4/t= 3EtchantChromic Acid: C = 300, 325, 350, 375 g/L and t = 2, 5, 10 min KMnO4: C = 40, 50, 60, 70 g/L and t = 5, 10, 15 minFigure 7.

UTS vs Etchant Type, Composition and Time for VeroBlueC = 0/t= 0C =1/t = 118.00C = 1/t =216.00C = 1/t= 3n o14.00C = 2/t= 1i t a g n o12.0010.00C = 2/t= 2C = 2/t= 3lE %8.00C = 3/t= 1C = 3/t= 26.00C = 3/t= 34.00C = 4/t= 1KMnO4Chromic AcidC = 4/t= 2EtchantC = 4/t= 3Chromic Acid: C = 300, 325, 350, 375 g/L and t = 2, 5, 10 min KMnO4: C = 40, 50, 60, 70 g/L and t = 5, 10, 15 minFigure 10. % Elongation vs Etchant Type, Composition and Time for VeroBluethe UTS decreased from 70 to 45 MPa with increase in etchant concentration, compared to 78 MPa for the base material (Fig. 7). This resulted in an average decrease in strength of 30%. The UTS for VeroBlue ranged from 35-39 MPa, compared to 42 MPa for the base material (Fig. 9). The % elongation decreased significantly for both materials and ranged between 1.39-2.15% for NanoForm, and between 6.53-9.48% for VeroBlue, compared to 2.36 and 17.46%, respectively, for the base materials.

This is not surprising, since at most concentrations and etch times, the potassium permanganate etchant resulted in blistering of NanoForm and severe chemical attack on VeroBlue.5. ConclusionsThis work explored the effect of the etching process on the mechanical properties of NanoForm and VeroBlue. Chromic acid and potassium permanganate were investigated at various concentrations and etch times as part of an electroless Ni process. The following bulleted list summarizes the important conclusions from this work.The chromic acid etchant resulted in a more uniform alteration of the surface for both NanoForm and VeroBlue compared to the potassium permanganate etchant; severe surface damage was observed with the potassium permanganate. The chromic acid etchant created uniform etch pits throughout the surface at most concentrations.Surface cracks were observed for etch times of 10 minutes for all chromic acid concentrations on the NanoForm material.

These cracks should be avoided in order to produce high quality films.Blistering of the NanoForm samples was observed for most concentrations of the potassium permanganate etchant.The chromic acid and potassium permanganate etchants resulted in a decrease in UTS for all concentrations and etch times for both NanoForm and VeroBlue.The degree of surface damage caused by the potassium permanganate etchant was severe on both NanoForm and VeroBlue, with an average decrease in UTS of 30% for NanoForm and 10% for VeroBlue. The % elongation decreased significantly for VeroBlue compared to the non-etched material.A smooth surface with fine surface modification throughout will result in the best electroless Ni coatings with a minimum effect on mechanical properties.The surface preparation for NanoForm consisting of chromic acid resulted in the highest quality Ni films for all concentration values.AcknowledgementsThe facilities within the W.M.

Keck Center for 3D Innovation (Keck Center) used here contain equipment purchased through Grant Number 11804 from the W.M. Keck Foundation, a faculty STARS Award from the University of Texas System, and two equipment grants from Sandia National Laboratories. This material is based in part upon work supported through the Mr. and Mrs. MacIntosh Murchison Chair I in Engineering, research contract 504004 from Sandia National Laboratories in the Laboratory Directed Research and Development (LDRD) program, and the National Science Foundation under Grant No. 0245071. The authors are grateful for the assistance of Mahesh Tonde, Karina Puebla, Luis Ochoa, Jesus Flores, Marjorie Ingle and Donovan Barroso within the Keck Center during performance of various aspects of the research.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of surface state and material on surface quality enhancement by Dual Laser Powder Bed FusionD. Ordnung*, J. Metelkova, and B. Van HoorewederDepartment of Mechanical Engineering, KU Leuven, 3001 Leuven, BelgiumAbstractParts produced by Laser Powder Bed Fusion typically exhibit a limited surface quality often requiring systematic post-processing. The KU Leuven AM team recently developed a Dual Laser Powder Bed Fusion strategy to improve the quality of inclined up-facing surfaces during building. It consists of two steps: (1) a pulsed laser induces shock waves to remove powder from inclined surfaces, followed by (2) in-situ laser remelting of the newly exposed surfaces. The first part of this paper covers the powder removal efficiency using shock waves depending on the used material.

The first part of this paper covers the powder removal efficiency using shock waves depending on the used material. A design of experiments was performed for horizontal samples of tool steels, titanium and aluminium alloys. The second part deals with the effect of the initial surface state on the powder removal efficiency for inclined surfaces (SaR,LT60=16.2 µm, SaR,LT120=24.0 µm). Finally, the third part demonstrates the surface quality improvement, resulting in a reduction of Ra up to 61% for 15° inclinations.IntroductionMetal parts produced by Laser Powder Bed Fusion (LPBF) typically exhibit a poor surface quality compared to conventionally manufactured parts. In particular, inclined up-facing surfaces pose a significant challenge as they are affected by so-called staircase and edge effects [1, 2]. Therefore, these surfaces usually require additional post-processing (e.g. machining) since they are inaccessible for surface treatment during LPBF.

Therefore, these surfaces usually require additional post-processing (e.g. machining) since they are inaccessible for surface treatment during LPBF. To address this issue, a Dual Laser Powder Bed Fusion (DLPBF) technique has been recently introduced to directly improve the quality of inclined up-facing surfaces during LPBF [3]. It is achieved with a modified commercial single-scanner LPBF setup enabling a consecutive operation of a continuous wave IR laser and a nanosecond pulsed IR laser, both working at 1060 nm wavelength. The technique consists of two steps: (1) using the pulsed laser to locally and selectively create shock waves which remove powder that is covering inclined surfaces, followed by (2) in-situ laser remelting of the newly exposed surfaces.Powder removal via shock waves is based on laser ablation, a technique that uses high intensity laser pulses to vaporize material.

Laser pulses interact with the solid metal surface, which rapidly heats up and vaporizes if the vaporization temperature is exceeded. The generated metal vapor plume is ionized upon interacting with subsequent laser pulses, creating a plasma plume [4]. The rapid expansion of the plasma plume induces radially propagating pressure waves known as laser-induced shock waves. The onset and strength of these shock waves is highly dependent on the specific material properties of the target such as absorptivity, density, thermal conductivity, melting and vaporization thresholds. The phenomenon of laser-induced shock waves is described in detail in [5, 6].Laser-induced shock waves are typically an undesired process by-product in LPBF, as they can blow away powder particles and molten metal preventing the formation of a stable track [7].

The authors of this paper exploited this phenomenon to remove powder from inclined surfaces, enabling subsequent improvement of the newly exposed surface via in-situ laser remelting. Previous studies on maraging steel M300 investigated various process parameters such as pulse energy, pulse-to-pulse distance, pulse duration and scan strategy, on the efficiency of powder removal [3, 8].Although the previous studies [3, 8, 9] have demonstrated a significant quality improvement of inclined up- facing surfaces, the efficiency of this technique applied to other AM materials remains unexplored, as does the effect of the initial surface roughness. This study showcases the feasibility of the DLPBF technique for treating various AM materials (M300, M789, A205, Ti64), based on a design of experiments including several pulsed laser parameters (pulse-to-pulse distance, pulse duration, pulse energy).

Then, the efficiency of powder removal from 15° inclined surfaces is assessed, considering M789 parts with variable initial surface state (built with a standard and increased layer thickness). Finally, in-situ quality improvement of the aforementioned M789 parts with inclined surfaces is demonstrated by combining selective powder removal and laser remelting during LPBF.380MethodologyPowder removal process study on horizontal surfacesTo evaluate the disturbance of the powder bed generated by shock waves, Fig. 1a shows a centered sample of 6 x 6 x 4 mm built by LPBF from different materials (M300, M789, A205, Ti64). For details on those four widely used materials used in AM and for the corresponding LPBF process parameter optimization the reader is referred to [10, 11, 12, 13] respectively.

For details on those four widely used materials used in AM and for the corresponding LPBF process parameter optimization the reader is referred to [10, 11, 12, 13] respectively. After building the last layer, the horizontal surfaces of the parts were scanned with variable pulsed laser parameters (Table 1) based on a Taguchi L9 design of experiments [8], inducing a disturbance of the surrounding powder bed. The samples were surrounded by a frame of 26 x 26 x 4 mm (Fig. 1a). In order to quantify the powder bed disturbance, the area inside the frame was scanned with the continuous wave (CW) laser (see fixation parameters in Table 2), gently melting the loose powder and hence fixing its shape. The disturbed powder bed was examined over an area of 1.3 x 25.4 mm (Fig. 1d) using a confocal focus variation microscope Sensofar S-Neox. The measured 3D data were filtered using a Gaussian filter, removing wavelengths < 250 μm.

1d) using a confocal focus variation microscope Sensofar S-Neox. The measured 3D data were filtered using a Gaussian filter, removing wavelengths < 250 μm. Following the analysis method described in detail in [8], the inner length of the disturbed powder bed was defined based on the calculated mean profile. The results show the mean inner disturbance length Li based on Li(1) and Li(2).Powder removal applied to inclined surfacesThe powder removal efficiency on inclined surfaces was assessed on 15° inclined M789 parts of variable initial surface state, resulting from using a standard or extremely high layer thickness of 60 µm and 120 µm (LT60 and LT120), respectively. As illustrated in Fig. 1b, after building the last layer, the pulsed laser was used to remove the powder from the inclined surfaces, while varying the number of scan passes (Table 1).

As illustrated in Fig. 1b, after building the last layer, the pulsed laser was used to remove the powder from the inclined surfaces, while varying the number of scan passes (Table 1). After the powder removal step, the inclined surface area was scanned with the CW laser in order to fix the partially removed powder (Table 2). The amount of removed powder was quantified by the mean cleaned length Lc based on the measured maximum and minimum cleaned length Lc(max) and Lc(min), respectively, as defined in Fig. 1b. Lc was then compared to the full part length Lref. Images of the cleaned surfaces were captured with a 3D digital microscope Keyence VHX-6000.Surface quality improvement by powder removal and subsequent laser remeltingQuality improvement was demonstrated on M789 parts with 15° inclined surfaces for LT60 and LT120 parts (Fig. 1c).

1c). Similarly to the previous experiment on inclined surfaces, after building the entire sample, the powder was removed from the inclined surface using pulsed laser and remelted using CW laser in order to improve the surface quality. Powder removal was done by scanning along the surface slope with the pulsed laser, starting from the accessible horizontal top surface towards the covered end of the inclination. The parameters are based on findings from [3, 8] and given in Table 1, except for the number of scan passes N = 15 and N = 20 for LT60 and LT120, respectively. Subsequent laser remelting by CW laser was performed by four consecutive scan passes under 45° with incremental rotation of 90° between passes.

Subsequent laser remelting by CW laser was performed by four consecutive scan passes under 45° with incremental rotation of 90° between passes. In addition to the parameters from Table 2, the focal plane was shifted towards the middle height of the inclined surface by moving the build cylinder up by f = 2 mmf = 1.3 mm for powder removal and laser remelting, respectively.𝛥𝛥The surface quality was assessed using a confocal focus variation microscope Sensofar S-Neox over an and 𝛥𝛥 area of 8.7 x 4.2 mm and 4.7 x 4.7 mm for inclined and horizontal surfaces, respectively.

Process parameters for fixation of the powder bed and laser remelting using CW laser. Process step Material Sample geometry (surface orientation) Laser power, P [W] Scanning speed, v [mm/s]Powder fixationM789 Cubic (horizontal)/ Prism (15° inclined) 100 1500 1300 40 2M300A205Ti64Cubic (horizontal)100 1500 1300 40 2100 1500 1300 40 2100 1500 1300 40 2Hatch spacing, h [μm] Number of passes, NLaser remelting M789Prism (15° inclined)20050080 4382Results & Discussion1. Powder removal using shock waves applied to various materialsIn the first part of this study, the effect of the used material on the powder removal efficiency was investigated. Fig. 2 shows the main effects plot based on a L9 Taguchi design of experiments for the inner disturbance length Li, which quantifies the disturbed powder bed around the parts. As shown in Fig.

Fig. 2 shows the main effects plot based on a L9 Taguchi design of experiments for the inner disturbance length Li, which quantifies the disturbed powder bed around the parts. As shown in Fig. 2a-c, Li exhibits similar trends for all investigated materials (M300, M789, A205, Ti64) within the selected parameter range. In agreement with the previously published results [8], the pulse duration τ has the most dominant impact on Li, followed by pulse energy Ep and pulse-to-pulse distance dp. Further on, Li and the generated shock wave efficiency appears to be higher with increasing pulse duration (Fig. 2c), pulse energy (Fig. 2b), and decreasing pulse-to-pulse distance (Fig. 2a). This becomes particularly evident for the pulse energy, as a larger amount of energy is deposited into the material, increasing the volume of interacting material. Consequently, a larger amount of material becomes available for vaporization and plasma formation, subsequently inducing the shock wave.

Consequently, a larger amount of material becomes available for vaporization and plasma formation, subsequently inducing the shock wave. The influence of the pulse-to-pulse distance on Li might be partially explained by the number of pulses available for inducing shock waves. Considering a constant scanned length and area, as the pulse-to-pulse distance decreases, the number of pulses increases and so does the number of shock waves. This is in accordance with the point- strong-explosion theory [6], which describes pulses as the initiators of local explosions that cause shock waves. The influence of pulse duration τ on Li is less evident. Studies [14] have reported contrary findings to the present research, i.e. an increase in shock wave pressure with decreasing τ, as shorter τ lead to a higher peak power. However, a previous research on the used equipment [15] reported a non-linear dependence of the efficiency of the laser induced shock waves on τ (within the range 10-100 ns).

However, a previous research on the used equipment [15] reported a non-linear dependence of the efficiency of the laser induced shock waves on τ (within the range 10-100 ns). This research confirmed a positive effect of increasing τ on the shock wave efficiency between 10-50 ns. This observation might be attributed to several phenomena associated with an extended laser-material interaction time and volume. Longer τ allow for interaction with a larger volume of material, potentially available for plasma formation and shock wave generation. This is supported by [8, 16] which reported a decrease of ablated material volume towards shorter τ in the same regime. Moreover, with an extended laser-material interaction time, the plasma plume has more time to form and expand before the exposure ends. This prolonged expansion could potentially enhance the strength of the shock wave, leading to an increased Li.

This prolonged expansion could potentially enhance the strength of the shock wave, leading to an increased Li. Nevertheless, as reported in [15], further increase in τ > 50 ns resulted in a progressively reducing shock wave efficiency. As a matter of fact, the peak power decreases with increasing τ, consequently reducing the amount of evaporated material turned into plasma plume.All investigated materials show similar trends, suggesting the occurrence of similar phenomena associated with the creation of laser-induced shock waves. However, there are variations in the magnitude of Li. Compared to the tool steels (M789, M300), the mean Li value for Ti64 is increased by 53.2 % and 32.5 %, respectively (Fig. 2d). The shock wave efficiency when processing A205 is comparable to Ti64, only 8.1 % lower.

2d). The shock wave efficiency when processing A205 is comparable to Ti64, only 8.1 % lower. To understand the differences in mean Li for the investigated materials, following aspects can be considered: (1) creation of shock waves, (2) displacement of the powder particles.The creation of laser-induced shock waves is well-known for metals and relates to various material properties, including absorptivity, thermal conductivity, melting and vaporization temperature. In general, metals with higher absorptivity facilitate more efficient coupling of the pulse energy, and hence for heating. A low thermal conductivity promotes localized heating, which is beneficial for reaching the vaporization temperature.

A low thermal conductivity promotes localized heating, which is beneficial for reaching the vaporization temperature. On the other hand, an increased thermal conductivity expands the heat-affected zone, increasing the volume of material potentially being available for material removal and thus the creation of shock waves.A205 and Ti64 possess distinctively different material properties, yet both materials lead to a comparable shock wave efficiency. Ti64 exhibits a low thermal conductivity and high melting and vaporization temperature, on the other hand A205 has a high thermal conductivity and low melting and vaporization temperature. Hence, reaching the vaporization temperature in Ti64 should require a high pulse energy input. However, as the thermal conductivity is low, the metal is heated up rather locally and should be hence available for creating strong shock waves.

However, as the thermal conductivity is low, the metal is heated up rather locally and should be hence available for creating strong shock waves. In case of A205, reaching those threshold temperatures required less pulse energy, which possibly leads to an early onset of vaporization. As the melting and vaporization temperatures are comparably low, and the thermal conductivity high, it should promote an extended heat-affected zone with potentially an increased amount of material available for vaporization. On the other hand, compared to A205 and Ti64, the investigated steels have an intermediate level of thermal conductivity and melting and vaporization temperatures. Yet, the effect of shock waves on M300 and M789 is weaker compared to the lighter A205 and Ti64. Therefore, it cannot be383excluded that the generated shock waves are of a comparable magnitude for all the investigated materials. Hence, material density is another important factor to be considered.

Hence, material density is another important factor to be considered. The low mass of the powder particles is expected to promote the disturbance of the powder bed, as they can be more easily displaced, which directly leads to an increase in Li.After investigating the effect of the used material on the shock wave efficiency, also the influence of extremely high roughness was studied. Such a roughness can be resulting from a high-productivity LPBF parameter set using a high layer thickness, such as reported in [13]. The Li observed for highly rough M789 surfaces generated by LT 120 µm was not included in Fig. 2 as it appeared to be very small, hindering a proper detection with this methodology.

The Li observed for highly rough M789 surfaces generated by LT 120 µm was not included in Fig. 2 as it appeared to be very small, hindering a proper detection with this methodology. Therefore, instead of studying the effect of the surface roughness on the powder removal efficiency applied to horizontal surfaces, the next section presents its application to inclined surfaces.Figure 2: Main effects plot for inner powder bed disturbance Li as a function of a) pulse-to-pulse distance, b) pulse energy, c) pulse duration, and d) mean inner disturbance length per material based on 9 samples each.2. Powder removal efficiency on inclined surfacesThis section described the powder removal efficiency on 15° inclined surfaces from M789, produced with a standard (LT60) and a high layer thickness (LT120). Although using a high layer thickness can increase the LPBF process productivity up to 91% [13], LT120 surfaces exhibit significantly higher roughness compared to LT60.

Although using a high layer thickness can increase the LPBF process productivity up to 91% [13], LT120 surfaces exhibit significantly higher roughness compared to LT60. The roughness values for investigated samples are SaR,LT60=16.2 µm and SaR,LT120=24.0 µm, corresponding to an increase by 48.2%.Fig. 3 shows the effect of number of scans (in powder removal) on the mean cleaned length Lc, serving as a measure of powder removal efficiency for parts with 15° inclined surfaces. Lc increases with the number of scan passes within the investigated range for both LT60 and LT120. Notably, LT60 exhibits a higher Lc for each investigated level of scan passes, up to 26.6% for N=1. However, the difference in powder removal efficiency for LT60 and LT120 appears lower with increasing scan passes, declining from 26.6% to 7.7% for N=1 and N=5 scan passes. Fig. 3 also shows that the difference between Lc(max) and Lc(min) reduces with increasing N for both LT60 and LT120.

Fig. 3 also shows that the difference between Lc(max) and Lc(min) reduces with increasing N for both LT60 and LT120. Although after N=1, a significant amount of powder is removed (high Lc(max)), low Lc(min) values indicate only a small fully powder-free area. Fig. 4 visualizes an asymmetric (Lc(max)>Lc(min)) and an approximately rectangular cleaned area (Lc(max)≈Lc(min)) for N=1 and N=5, respectively. A previous study [3] observed a similar phenomenon, attributing it to a redeposition of powder on the already cleaned surface. Due to the large powder volume covering the inclined surface during the first scan pass, the powder partially falls back on the already powder-free surface in the wake of the laser. With subsequent scan passes, the powder volume on the inclined surface decreases, and so does the redeposition of powder, resulting in an approximately rectangular cleaned area (Lc(max)≈Lc(min)).

With subsequent scan passes, the powder volume on the inclined surface decreases, and so does the redeposition of powder, resulting in an approximately rectangular cleaned area (Lc(max)≈Lc(min)). Recent research has shown that powder can be entirely removed from inclined surfaces by further increasing the number of scan passes [3, 8, 9].Powder removal via laser-induced shock-waves can be regarded as a surface phenomenon, given that both LT60 and LT120 were manufactured from the same powder material. As stated above, LT120 samples exhibit a significantly higher surface roughness. Consequently, the significant increased Lc for LT60 with N=1 scan pass can be attributed to the lower surface roughness. From the second powder removal pass, the process happens on a modified surface. The effect of powder removal on the surface was observed in [9], showing a significant increase of the overall surface texture for inclined surfaces after repeated powder removal. Fig.

The effect of powder removal on the surface was observed in [9], showing a significant increase of the overall surface texture for inclined surfaces after repeated powder removal. Fig. 3 shows that the amount of removed material per pass reduces with each following pass. Hence, this efficiency reduction could be attributed to the gradual deterioration of the initial high surface quality with each scan pass. The presence of large384roughness peaks and valleys could potentially dissipate energy, thereby reducing the available energy for shock wave generation. Additionally, the increased size of these surface features might dampen the propagation of shock waves and the motion of powder particles.

Additionally, the increased size of these surface features might dampen the propagation of shock waves and the motion of powder particles. A similar reasoning can be made for a decreased Lc on rough LT120 surfaces.Figure 3: Mean cleaned length Lc given in % of absolute length of the inclined surface Lc(ref) as function of number of scan passes N, with the error bars indicating mean maximum and mean minimum cleaned length Lc(max), and Lc(min), respectively. Reported mean values are based on 6 samples each.Figure 4: Top view on inclined surfaces of M789 samples after powder removal, a) LT60 surface after N=1, and LT120 surface after b) N=1 and c) N=5.3. Proof of concept - Surface quality improvementThe proposed method has been successfully applied to improve the quality of inclined surfaces from M789, produced with a standard (LT60) and a high layer thickness (LT120).

The powder covering the 15° inclined surface could be entirely removed via laser-induced shock waves, after N=15 and N=20 scan passes for LT60 and LT120, respectively. This is in agreement with the findings from above, showing an increased powder removal efficiency for LT60 surfaces. By remelting of the inclined surface the initial as-built surface roughness could be reduced from SaR,LT60=16.2 µm to 6.4 µm and SaR,LT120=24.0 µm to 16.4 µm for LT60 and LT120. This corresponds to a significant improvement by 61% and 32%, respectively. Surface quality improvement in this order of magnitude on inclined surfaces from M789 produced with LT60 has been previously reported [17]. Although laser remelting demonstrates a notable reduction in surface roughness for investigated surfaces, the effect of LT120 is more moderate. This can be attributed to the fact that laser remelting parameters were optimized for LT60 surfaces.

This can be attributed to the fact that laser remelting parameters were optimized for LT60 surfaces. Moreover the initial variation in surface roughness impacts the efficiency of laser remelting. Laser remelting involves redistributing of molten material from roughness peaks to valleys. As LT120 surfaces exhibit larger roughness peaks and valleys, the molten material needs to fill equally larger cavities.385ConclusionThis work demonstrates that the innovative Dual Laser Powder Bed Fusion strategy allows treating various AM metals during building (M300, M789, A205, Ti64), despite very different material properties. This technique combines selective powder removal via laser-induced shock waves and laser remelting during building. The powder removal efficiency on horizontal surfaces followed similar trends for all investigated materials (M300, M789, Ti64, A205), but the effect appeared to be stronger for lighter materials.

The powder removal efficiency on horizontal surfaces followed similar trends for all investigated materials (M300, M789, Ti64, A205), but the effect appeared to be stronger for lighter materials. It decreased with pulse- to-pulse distance, and increased with pulse energy and pulse duration. The pulse duration was found to have a dominant impact. The powder removal efficiency on inclined surfaces was examined on M789 parts inclined by 15°, produced with a standard (LT = 60 µm) and a high layer thickness (LT = 120 µm). The powder removal efficiency appeared to be higher on surfaces produced with LT60, which was attributed to the significantly lower surface roughness compared to LT120. However, the difference in powder removal efficiency for both LT appeared to decrease with the number of scan passes, suggesting a LT60 surface roughness deterioration due to the powder removal.

However, the difference in powder removal efficiency for both LT appeared to decrease with the number of scan passes, suggesting a LT60 surface roughness deterioration due to the powder removal. Finally follows a demonstration of the quality improvement of 15° inclined surfaces from M789 parts by Dual Laser Powder Bed Fusion. After laser remelting a significant surface quality improvement was achieved, reducing SaR by 61% and 32% for LT60 and LT120, respectively.The presented Dual Laser Powder Bed Fusion strategy allows treating various AM metals during building. The improvement of inclined surfaces manufactured with high layer thickness offers a significant value for additive manufacturing as it combines two typical contrary aspects: high productivity and improved surface quality.

Moreover, as this treatment happens during building, it has the potential to reduce or eliminate the need of additional post-machining steps.AcknowledgementsThis research was funded by VLAIO/SIM/ICON HBC.2020.2958 MetaMould and by FWO/SB1SB2322N. The authors would like to thank 3DSystems for providing certified building parameters for M789.References [1] J. Metelkova, L. Vanmunster, H. Haitjema, and B. Van Hooreweder, “Texture of inclined up-facing surfaces fusion of metals,” Addit. Manuf., p. 101970, Mar. 2021, doi:in laser powder bed 10.1016/j.addma.2021.101970.[2] G. Strano, L. Hao, R. M. Everson, and K. E. Evans, “Surface roughness analysis, modelling and prediction in selective laser melting,” J. Mater. Process. Technol., vol. 213, no. 4, pp. 589–597, Apr. 2013, doi: 10.1016/j.jmatprotec.2012.11.011.[3] J. Metelkova, D. Ordnung, Y. Kinds, and B.

fect of Surface Treatments on Electron Beam Freeform Fabricated Aluminum StructuresKaren M. B. Taminger, Robert A. Hafley, David T. Fahringer, and Richard E. Martin NASA Langley Research Center, Hampton, VAReviewed, accepted August 4, 2004AbstractElectron beam freeform fabrication (EBF3) parts exhibit a ridged surface finish typical of many layer-additive processes. Thus, post-processing is required to produce a net shape with a smooth surface finish. High speed milling, wire electrical discharge machining (EDM), electron beam glazing, and glass bead blasting were performed on EBF3-built 2219 aluminum alloy parts to reduce or eliminate the ridged surface features. Surface roughness, surface residual stress state, and microstructural characteristics were examined for each of the different surface treatments to assess the quality and effect of the surface treatments on the underlying material.

The analysis evaluated the effectiveness of the different surface finishing techniques for achieving a smooth surface finish on an electron beam freeform fabricated part.IntroductionThe electron beam freeform fabrication (EBF3) process has been developed over the past two years at NASA Langley Research Center. A computer model of a component is created and translated into programmable machine code for driving the EBF3 system. A focused electron beam is used to create a molten pool on a metal substrate. Wire is fed into the molten pool and the substrate is translated with respect to the electron beam and wire to build a layer. The final part is, thus, built up layer by layer. However, as with many layer-additive processes, the surface exhibits a ridged surface finish corresponding to the deposited layers.

The final part is, thus, built up layer by layer. However, as with many layer-additive processes, the surface exhibits a ridged surface finish corresponding to the deposited layers. If the component is designed to be built as a near-net shaped part in the freeform fabrication system, then some form of post-processing is planned as part of the manufacturing life cycle to achieve final dimensional tolerance and adequate surface finish for the intended application. In this case, as long as the surface finish is conducive to performing secondary machining, (i.e. the ridges in the surface do not induce chatter or otherwise impact the performance of the tool), then the surface finish is not a significant issue. However, for other applications, it is desirable to achieve final dimensions and surface finish directly from the fabrication process.

However, for other applications, it is desirable to achieve final dimensions and surface finish directly from the fabrication process. To accomplish production of net-shaped parts directly in the EBF3 system, techniques need to be developed to produce an acceptable surface finish.Experimental ProceduresThe purpose of this work was to compare different surface finishing techniques forsmoothing the ridged layer-additive surface finish of parts fabricated using the EBF3 process. Five pylons were fabricated from 0.063 in. diameter aluminum alloy 2219 (Al-6% Cu nominal composition) wire using the EBF3 process. The pylons were each 4.5 in. long, 1.5 in. wide, and 2 in. high with a wall thickness of approximately 0.25 in. A pylon in the as-built condition is460shown in Figure 1. The pylons were used to demonstrate finishing techniques on a contoured component. A straight wall was also built 10 in. long, 2 in. high, and approximately 0.25 in. thick.

The pylons were used to demonstrate finishing techniques on a contoured component. A straight wall was also built 10 in. long, 2 in. high, and approximately 0.25 in. thick. This straight wall was cut into five 2-in. long segments, which were subjected to the same finishing treatments as the pylons. These wall segments were used for residual stress and surface finish measurements and then were sectioned for metallurgical analysis. Since all of the wall segments came from the same build, comparing data on these straight wall segments was a direct measure of the surface treatments and was not influenced by any inconsistencies that may occur in the EBF3 fabrication process from part to part. The deposition rate for the pylons and the straight wall was approximately 12 in.3/hr, resulting in a build time of 6 hr. per pylon.1 in.Figure 1.

The deposition rate for the pylons and the straight wall was approximately 12 in.3/hr, resulting in a build time of 6 hr. per pylon.1 in.Figure 1. Electron beam freeform fabricated 2219 Al pylon in as-built condition.The surface treatments evaluated were: high speed milling, wire electrical dischargemachining (EDM), glass bead blasting, and electron beam glazing. A brief description of the conditions used for performing these surface finishing operations follows. These surface treatments were compared to the as-built condition based on an evaluation of the surface roughness and waviness, surface residual stress state imposed into the part, and the impact of the surface treatment on the near-surface microstructure.High Speed MillingFigure 2 shows a pylon after highspeed milling on the inside and outside surfaces. A 0.500-in. diameter, 4 flute, high speed steel end mill was used to remove approximately 0.01-0.02 in. from each surface at a feed rate of 15 in./min.

A 0.500-in. diameter, 4 flute, high speed steel end mill was used to remove approximately 0.01-0.02 in. from each surface at a feed rate of 15 in./min. The machining operation consisted of two passes, a roughing pass and a finishing pass, and took approximately 4 min. of actual machining time. The pylon solid model developed for the EBF3 system was also used to program the machining, so minimal time was required to program the milling operation. However, additional time was spent: (1) removing the part from the EBF3 system, (2) setting up the pylon in the high speed milling machine, and (3) positioning the near-net shape part with respect to the machining file to perform the machining. Similar machining set-up and feed rates were also used to cut the 2-in. straight wall segment down the centerline of the deposit for analysis. The operator reported that the machinability of the EBF3 material was comparable to typical aluminum.1 in.Figure 2.

straight wall segment down the centerline of the deposit for analysis. The operator reported that the machinability of the EBF3 material was comparable to typical aluminum.1 in.Figure 2. Electron beam freeform fabricated 2219Al pylon after high speed milling.461Wire EDMFigure 3 shows a pylon after beingseparated from the baseplate and wire EDM on the inside and outside surfaces. A 0.010 in. diameter wire was used to EDM the pylon. Actual wire EDM time was approximately 10 min. to remove approximately 0.05 in. from the inside and outside of the pylon. Prior to machining the pylon surfaces, a pilot hole was drilled into the baseplate in the center of the pylon because the wire must pass through the part to carry the current during wire EDM, preventing wire EDM from being effective for plunge operations. In addition, supports were left on either side near the center of the pylon to avoid pinching the wire.

In addition, supports were left on either side near the center of the pylon to avoid pinching the wire. (The remnants of these supports are still visible as vertical stripes in Figure 3.) These supports were then removed in a subsequent machining operation. As with the high speed milled pylon, the pylon solid model developed for the EBF3 system was also used to program the wire EDM system, minimizing programming time. However, additional time was spent: (1) removing the part from the EBF3 system, (2) setting up the pylon in the wire EDM, (3) positioning the near-net shape part with respect to the machining file to perform the machining, and (4) final removal of the supports. Although wire EDM is typically followed by a grinding operation to remove the rough recast layer, the recast layer was not removed from the pylon to enable surface finishing analysis. Similar power settings and feed rates were also used to cut the 2-in.

Similar power settings and feed rates were also used to cut the 2-in. straight wall segment down the centerline of the deposit for analysis. Aside from the additional steps required for the wire EDM process, the operator reported no problems with machining the parts fabricated using the EBF3 process.1 in.Figure 3. Electron beam freeform fabricated 2219Al pylon after wire EDM machining.Glass Bead BlastingFigure 4 shows a pylon after glassbead blasting the exterior surface of the pylon. The pylon was bead blasted for 4 min. at a feed rate of approximately 5 in./min. using 0.015 in. diameter glass bead media. The set-up time was minimal since the bead blasting was manually controlled. Due to the geometry of the pylon, it was not possible to reach the inside surfaces of the pylon effectively. The 2-in. straight wall segment was bead blasted on both surfaces using feed rates and times similar to those used for the pylon.1 in.Figure 4.

The 2-in. straight wall segment was bead blasted on both surfaces using feed rates and times similar to those used for the pylon.1 in.Figure 4. Electron beam freeform fabricated 2219Al pylon after glass bead blasting.462Electron Beam GlazingFigure 5 shows a pylon that was electron beam glazed in the EBF3 system immediately after completing the fabrication. The pylon was tilted 90° from the build orientation, and a low power, defocused electron beam was rastered perpendicular to the build direction to remove the EBF3 surface features. Total processing time was approximately 5 min., but no additional set-up time was required to perform the glazing operation. However, only the external surface of the pylon was accessible to the electron beam during the glazing operation, so it was not possible to reach the inside surfaces of the pylon effectively. Since the straight wall had to be removed from the EBF3 system to be cut into 2-in. segments, the 2-in.

Since the straight wall had to be removed from the EBF3 system to be cut into 2-in. segments, the 2-in. segment was returned to the EBF3 system and electron beam glazed using beam power and rastering rate and pattern similar to those used for the pylon.Build DirectionGlaze Direction1 in.Figure 5. Electron beam freeform fabricated 2219Al pylon after electron beam glazing.Microstructural AnalysisCross sections of the 2-in. straight wall segments were polished and etched with Keller’s reagent. The near-surface microstructures of the wall segments subjected to surface treatments were compared to that of the 2-in. straight wall segment in the as-built condition. The segments were photographed using an optical metallograph to examine the impact of the surface treatments on the microstructure of the segments.Surface Roughness and WavinessSurface roughness and waviness measurements were made on the 2-in.

straight wallsegments for each surface treatment investigated, and compared to the as-built condition. Surface roughness measurements were taken using a laser interferometer with a diamond stylus. The sensitivity of the measurements on the instrument used was +/- 0.1 µin. Measurements of the root mean square (RMS) surface finish were taken over a distance of 0.09 in., measured across the layers (normal to the baseplate on which the part was fabricated) at a location in the center of the 2-in. build height. However, since a standard RMS finish measurement is highly localized, the surface waviness was also measured. Surface waviness measurements were made using the same laser interferometer over a 1.0 in. length across the layers, starting approximately 0.5 in. above the baseplate.Surface Residual StressThe surface residual stress state was measured on the center of the 2-in. straight wallsegments for each surface treatment investigated, and compared to the as-built condition.

straight wallsegments for each surface treatment investigated, and compared to the as-built condition. The x-463ray diffraction method for measuring residual stresses is appropriate for assessing the surface residual stress state imparted by the different surface treatments because it measures residual stresses in a very shallow volume of material. The d vs. sin2ψ technique [1] uses the displacement of atomic planes as an internal measurement of macro residual stresses near the surface of the specimen. The residual strain was determined by measuring the d-spacings of grains oriented at a variety of different angles with respect to the surface normal. The stress was then calculated from this measured residual strain using the bulk modulus of elasticity for 2219 Al. The experiments were performed using copper Kα radiation, in which the depth of penetration in the aluminum specimens was 0.4 µin. Measurements were taken over an area of approximately 0.25 in. by 0.25 in.

Measurements were taken over an area of approximately 0.25 in. by 0.25 in. square at a location in the center of each straight wall segment. Residual stresses were measured in three orientations: parallel to the build direction (designated 0º), perpendicular to the build direction (across the layers, designated 90º), and 45º to the build direction (designated 45º).Results and DiscussionLow and high magnification optical microstructures of the near-surface regions on each of the straight wall segments are shown in Figures 6-10 for the as-built, high speed milled, wire EDM, glass bead blasted, and electron beam glazed specimens respectively. In the as-built specimen, two deposit layers and the remelted zone between those two layers are evident in Figure 6a. The localized surface is quite smooth, but there is evidence of longer range periodic surface waviness corresponding to the layer height, which is approximately 0.03 in. for the parts fabricated in this study.

for the parts fabricated in this study. The surface microstructure shown in Figure 6b is typical of the through- thickness microstructure, showing relatively large grains with some evidence of a solidification dendrite structure within the grains. After high speed milling, Figure 7a shows scallops on the bottom of the end mill cut, but the sidewall of the machined segment is smooth with no evidence of surface variations. Figure 7b shows that the grain structure of the 2219 Al was not affected by the milling. After wire EDM, there is no longer any evidence of long range surface waviness in Figure 8a. Figure 8b shows that the recast layer is less than 200 µin. thick, but this layer causes asperities on a fine scale. After glass bead blasting, Figure 9a shows the long range waviness associated with the prior build layers remains intact. In addition, localized surface roughness in Figure 9b appears higher that that of the as-built segment in Figure 6b.

In addition, localized surface roughness in Figure 9b appears higher that that of the as-built segment in Figure 6b. Finally, after electron beam glazing, the surface appears very smooth with no long range waviness in Figure 10a. More noticeable is the fine-grained, equiaxed microstructure at a constant 0.04 in. depth associated with the heat affected zone from the glazing pass. The microstructure in the near-surface region after electron beam glazing (Figure 10b) is much finer than that of the as-built microstructure in Figure 6b.The micrographs qualitatively indicate the roughness of the as-built surface as comparedto that resulting from the various surface treatments. Surface roughness and waviness measurements were taken on the same straight wall segments as photographed in Figures 6-10. These results are shown in Table 1 and Figure 11.

Surface roughness and waviness measurements were taken on the same straight wall segments as photographed in Figures 6-10. These results are shown in Table 1 and Figure 11. Note that lower RMS values are associated with smoother surface finishes, with a mirror surface typically having an RMS value less than 8.464Layer “B” Layer “B”Build direction Build direction Build directionRemelted zone Remelted zone between layers between layers “A” & “B” “A” & “B”Layer “A” Layer “A”(a) (a) (a)(b) (b) (b)500 µm 500 µm 500 µm 500 µm50 µm50 µm 50 µm 50 µmFigure 6. Microstructure of as-built electron beam freeform fabricated straight wall segment showing (a)long range waviness, and (b) surface microstructural details.Machined Machined sidewall sidewallBottom of Bottom of end mill cut end mill cutBuild direction Build direction Build directionAs-built surface As-built surface at pylon base at pylon base(a) (a)(b) (b)500 µm 500 µm 500 µm50 µm50 µm 50 µmFigure 7.

Microstructure of electron beam freeform fabricated straight wall segment after glass beadblasting showing (a) long range waviness and (b) surface roughness.466Build direction Build direction Build directionDepth of heat Depth of heat affected zone affected zone during glazing during glazing pass pass(a) (a) (a)(b) (b) (b)500 µm 500 µm 500 µm 500 µm50 µm50 µm 50 µm 50 µmFigure 10. Microstructure of electron beam freeform fabricated straight wall segment after electron beamglazing showing (a) smooth surface with heat affected zone and (b) fine, equiaxed, near- surface microstructure.Table 1.

Surface roughness measurements for as-built and surface treated EBF3 deposits.Surface TreatmentRMS Roughness (µin.)Waviness (µin.)As-built9230,000High Speed Milled8-56400Wire EDM1961,500Bead Blasted23312,000Electron Beam Glazed184,000As the values for the as-built straight wall segment show in Table 1, it is possible toachieve a relatively low RMS surface finish yet have very high long range waviness. The as- built electron beam freeform fabricated segment has a reasonably low surface finish RMS value due to the smooth nature of the solidification of the molten pool. However, the layers result in a very high measure of waviness. Of the surface treatments studied, the electron beam glazed part produced the lowest RMS value, challenged only by high speed milling.

However, the layers result in a very high measure of waviness. Of the surface treatments studied, the electron beam glazed part produced the lowest RMS value, challenged only by high speed milling. The surface finish range shown for the milling machined specimen represents a range of surface finished achievable using this method – smaller step sizes can be applied to attain lower RMS values. Surface467asperities produced by the recast layer in the wire EDM machined segment and by the impact damage of the glass bead blasting actually increased the surface roughness RMS values versus the as-built segment.In Figure 11, all of the waviness plots are shown on the same scale, providing directcomparison of the surface conditions resulting from each of the different surface treatments.

All of the surface finishing techniques except for the glass bead blasting were successful at significantly reducing the surface waviness from the prior layers in the electron beam freeform fabricated segments.

The periodicity of the slight waviness that remained in the electron beam glazed segment is associated with the width of the raster pattern used to perform the glazing operation.30000 3000020000 2000010000 10000As-built EBF3 As-built EBF3 process process0 030000 3000020000 2000010000 10000High speed High speed machined machined0 0, , t t n n e e m m e e c c a a p p s s D Dl li is s e e h h c c n n o o r r c c m mi ii i30000 3000020000 2000010000 10000Wire EDM Wire EDM machined machined0 030000 3000020000 20000Glass bead blasted Glass bead blasted10000 100000 030000 3000020000 2000010000 10000Electron beam Electron beam glazed glazed0 00 00.2 0.20.4 0.40.6 0.60.8 0.81.0 1.0Scan Distance, in Scan Distance, inFigure 11.

Surface waviness measurements on straight wall segments for each surface treatment studied.Figure 12 shows the surface residual stress measurements from each straight wall segment parallel to the build direction (0º), perpendicular to the build direction (90º), and 45º to the build direction. The residual stresses are significant, but the magnitude of these values may not be accurate since the calculation of residual stresses is based upon the bulk tensile modulus468of the 2219 Al. There are two possible sources of error in the residual stress measurements shown in Figure 12. First, the modulus of the specific {hkl} planes of atoms measured at the atomic scale should be used to convert from residual strain to residual stress using the x-ray diffraction method. Since atomic-level moduli are difficult to determine and generally unavailable, it is standard practice to use a bulk modulus value as a best approximation. [1] However, this assumption may not be very accurate.

[1] However, this assumption may not be very accurate. Second, the yield and ultimate tensile strengths of as-deposited 2219 Al are approximately 17 ksi and 40 ksi respectively. [2] Thus, this conversion is invalid for stresses exceeding the yield strength of the material. Therefore, when examining the measured residual stresses, the relative values are more important than the specific magnitudes shown.The as-built segment exhibits moderate compressive residual stresses in the 0º orientationand low tensile residual stresses in the 45º and 90º orientations. These residual stresses are the result of thermal stresses, and are comparable to residual stresses expected from welding processes. [3] High speed milling, wire EDM, and glass bead blasting all induce compressive residual stresses in all directions. These measured results are consistent with handbook values for these finishing processes.

These measured results are consistent with handbook values for these finishing processes. [4] Electron beam glazing does not significantly change the compressive residual stress state in the 0º orientation or the tensile residual stresses in the 90º orientation versus the as-built segment. However, electron beam glazing does induce high compressive residual stresses in the 45º orientation as compared to the slightly tensile stresses measured in the as-built segment.100Residual Stress, ksi10203040As-builtMachined Wire EDMBead blastedGlazedSurface TreatmentFigure 12.

Surface residual stresses measured via x-ray diffraction on straight wall segments for eachsurface treatment studied.4690 45 90ConclusionsElectron beam glazing conducted in the EBF3 system is a viable process for removing surface waviness associated with EBF3 deposited layers, producing an RMS 18 surface finish with fine-grained equiaxed surface microstructure.The best combination of low localized RMS surface finish and minimal long range waviness was achieved using high speed milling machining.Flat and contoured electron beam freeform fabricated parts were successfully machined using high speed milling and wire EDM. Both processes induced compressive residual stresses into the parts as is typical for these processes.Glass bead blasting was not aggressive enough to eliminate the ridged surface finish resulting from the EBF3 process.References[1] Noyan, I. C. and Cohen, J.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of the Print Bed Temperature on Void Distribution within the Microstructure of Short Carbon Fiber Reinforced/ABS Manufactured via Large Area Additive ManufacturingNeshat Sayah, Dr. Douglas E. Smith, Department of Mechanical Engineering, Baylor University, Waco, TX, USAABSTRACTShort carbon fiber-reinforced polymer composite structures produced using Large Area Additive Manufacturing (LAAM) have garnered significant attention due to the design flexibility, energy savings, and materials selection associated with this process. However, the physical and mechanical properties of the additively manufactured composite parts often fall below expectations due to void formation between printed beads and within the microstructure of individual beads.

This study aims to investigate the effect of bed temperature on the microstructure within the beads of two-layer Short Carbon Fiber reinforced Acrylonitrile Butadiene Styrene (SCF/ABS) beads manufactured via the LAAM system. This study employs high-resolution 3D micro-computed tomography (µCT) to evaluate the void shape and distribution within the microstructure of composite parts printed at various bed temperatures. The results of this study demonstrate substantial variation in the void volume fraction among four bead sets deposited at different print bed temperatures. Moreover, within each part, a noticeable discrepancy in void volume fraction is observed between the top and bottom bead of the two-bead test samples. Preliminary results indicate that increasing the bed temperature from 25°C to 75°C reduces void volume fraction within the microstructure of the composite parts.

Preliminary results indicate that increasing the bed temperature from 25°C to 75°C reduces void volume fraction within the microstructure of the composite parts. However, an opposite trend emerges when the bed temperature is further increased to 100°C, increasing void volume fraction, which needs further investigation to understand. This study also evaluated the void shapes through the calculation of their sphericity. The preliminary results reveal that as the bed temperature increases from 25°C to 75°C, the voids exhibit higher sphericity within the printed parts as the interconnected voids decrease.INTRODUCTIONPolymer composite extrusion-deposition is a widely recognized and favored technique in Additive Manufacturing (AM). This technique stands out among the various AM methods due to the available range of neat thermoplastic polymers and thermoplastic polymer composites that can be utilized [1,2].

This technique stands out among the various AM methods due to the available range of neat thermoplastic polymers and thermoplastic polymer composites that can be utilized [1,2]. Polymer extrusion deposition AM can be implemented in two forms: small scale using Fused Filament Fabrication (FFF) with a liquefier or nozzle filament and large scale employing LAAM with a single screw extruder. This article focuses on the polymer composite extrusion-deposition process via LAAM [3]. Short carbon fibers are used as reinforcements in polymer composites due to improved mechanical and thermal properties compared to neat polymers. However, the properties of the final manufactured polymer composite parts are greatly influenced by their microstructure, which includes factors such as fiber orientation, voids between beads (inter-bead voids), and microstructural voids (intra-bead voids) [4-7].

Insufficient interlayer bonding during polymer composite deposition leads to inter-bead voids forming while intra-bead voids develop within the bead microstructure near the fibers [8-10]. Furthermore, microstructural226voids arise from air entrapment during the manufacturing process and from pre-existing voids present within the filaments or pellets before their utilization in the additive manufacturing process [7,11].This article specifically centers on microstructural intra-bead voids, which are named ‘micro voids’ within the body of this article. Optimizing the printing process parameters (i.e., nozzle temperature, bed temperature, nozzle height above the print surface, and extrusion rate) holds substantial potential for enhancing the quality and properties of the final 3D printed polymer composite [12-14]. Most prior research has considered the microstructure of deposited polymer composites employing two-dimensional (2D) imaging techniques, such as electron microscopy.

Most prior research has considered the microstructure of deposited polymer composites employing two-dimensional (2D) imaging techniques, such as electron microscopy. However, the sectional nature of this technique introduces a potential bias when evaluating voids, which can result in errors in the analysis outcomes [15-16]. More recently, researchers have increasingly employed µCT as a viable alternative to avoid potential errors associated with the sectional nature of 2D imaging techniques. µCT offers the advantage of generating high-resolution the three-dimensional (3D) structural microstructure of additively manufactured polymer composites [17-18]. Despite the previous research on extrusion deposition of polymer composites, there remains to be more information regarding the impact of 3D printing processing parameters on the microstructure of the final manufactured composites, particularly concerning microstructural voids.

This study aims to address this gap by presenting an experimental investigation focused on the influence of bed temperature on the void volume fraction, void distribution, and void shape within the microstructure of two-layer SCF/ABS composite beads 3D printed via LAAM, utilizing µCT as the imaging technique.information for various applications,includingMATERIALS AND METHODS3D Printing ProcessThis research uses the Baylor LAAM system to produce two-layer SCF/ABS polymer composite beads. The Baylor LAAM system appearing in Figure 1 is a custom-built gantry system that incorporates the Strangpresse Model 19 extruder (Strangpresse, Youngstown, OH, USA) [19], which features a single screw extruder with three temperature zones and a nozzle tip diameter of 3.172 mm. For the LAAM feedstock, 13 wt% short carbon fiber (having a diameter of 7.5 microns) PolyOne SCF/ABS pellets (Avient Corporation, Avon Lake, OH, USA) were used.

For the LAAM feedstock, 13 wt% short carbon fiber (having a diameter of 7.5 microns) PolyOne SCF/ABS pellets (Avient Corporation, Avon Lake, OH, USA) were used. Before starting the 3D printing process, the pellets were dried at 80°C for 12 hours in the oven.Figure 1. Baylor’s custom-built LAAM system [11]227This study used four different bed temperatures (25, 50, 75, and 100 °C) to 3D print two-layer SCF/ABS beads in order to evaluate the effect of bed temperature on their microstructure. It is worth noting that the remaining printing parameters, including the nozzle height above the print bed, nozzle temperature, and extrusion rate, remained unchanged across all printed components. The specific process parameters employed are outlined in Table1.Nozzle Height Nozzle Temperature Screw Speed Print Speed1.1 (mm) 200 (°C) 90 (rpm) 101.6 (mm/s)Table 1.

The specific process parameters employed are outlined in Table1.Nozzle Height Nozzle Temperature Screw Speed Print Speed1.1 (mm) 200 (°C) 90 (rpm) 101.6 (mm/s)Table 1. Printing process parametersµCT Image Acquisition TechniqueTo acquire accurate and detailed 3D microstructural information of the additively manufactured SCF/ABS parts, the NSI X3000 X-ray µCT system (North Star Imaging, Rogers, MN, USA) was employed in this study. All computed tomography (CT) images were captured at a resolution of 1.7 microns, while the X-ray source was set to generate a voltage of 60kV and a current of 750 µA. Each sample rotated 360 degrees, with increments of 2 degrees, resulting in 1800 projections. The raw data from the µCT system was subsequently imported into efX-CT software (North Star Imaging, Minnesota, USA) for 3D volume reconstruction. The reconstructed data was then imported into VGStudio Max 3.4 software (Volume Graphics GmbH, Heidelberg, Germany) for void analysis.

The reconstructed data was then imported into VGStudio Max 3.4 software (Volume Graphics GmbH, Heidelberg, Germany) for void analysis. The VGDefX algorithm was employed for void evaluation, which used grayscale data at the voxel scale to detect and analyze voids within the microstructure of each scanned part [11,20].RESULTS AND DISCUSSIONSAs discussed earlier in Section 2 of this study, µCT offers comprehensive microstructural details of the scanned parts. Figure 2 shows a 2D cross-section sample from a SCF/ABS bead printed on a bed with a temperature of 100°C which highlights the observed microstructural, including micro voids, short carbon fibers, and the polymer matrix.ABS MatrixMicro VoidShort Carbon FiberFigure 2. Microstructural features of SCF/ABS part printed on a bed with a temperature of 100°C2283 mm3 mm(a)(b)yy3 mm3 mmxx(c)(d)Figure 3.

Microstructural features of SCF/ABS part printed on a bed with a temperature of 100°C2283 mm3 mm(a)(b)yy3 mm3 mmxx(c)(d)Figure 3. 2D cross-sections obtained from µCT scans of the two-layer SCF/ABS beads manufactured on beds with different temperatures: a) print bed temperature of 25°C, b) print bed temperature of 50°C, c) print bed temperature of 75°C, and d) print bed temperature of 100°C.Figure 3 presents the 2D cross-sections of the four scanned two-layer SCF/ABS beads that were printed for this study. The printing parameters remained consistent across all samples, except for the bed temperature variation for each additively manufactured part, as explained in Table 1. The µCT images reveal variations in the bead's shape, void content, and void shape within the microstructure of each two-layer bead. Also shown is a significant variation in microstructure over the cross-section of the bead.

Also shown is a significant variation in microstructure over the cross-section of the bead. A significant presence of micro voids is evident within the bottom and top beads printed on a print bed temperature of 25°C, as shown in Figure 3(a). Furthermore, notable occurrences of large micro voids and interconnected voids are particularly prominent along the edges and near the print bed area. Figure 3(b) shows the microstructure of the two-layer bead printed on a print bed with a temperature of 50°C, revealing a significant number of micro voids. However, compared to the two-layer beads printed on a print bed temperature of 25°C (Figure 3(a)), the quantity of large micro voids and interconnected voids is reduced, particularly near the print bed area and the edges. However, it is essential to note that there is still evidence of large and interconnected voids near the center of both the top and bottom beads, as highlighted in Figure 3(b).

However, it is essential to note that there is still evidence of large and interconnected voids near the center of both the top and bottom beads, as highlighted in Figure 3(b). Figure 3(c) illustrates that the quantity of voids within the microstructure of the two-layer beads printed on a print bed with a temperature of 75°C has significantly decreased compared to those printed on print beds at 25°C and 50°C. Furthermore, micro voids appearing in Figure 3(c) are notably smaller, with very few large micro voids in both the top and bottom bead229microstructures. In the microstructure of the two-layer beads 3D printed on a print bed with a temperature of 100°C, a notable presence of voids is observed in both the top and bottom beads, as shown in Figure 3(d). The void content is relatively lower than in the bead printed on a print bed with a temperature of 75°C.

The void content is relatively lower than in the bead printed on a print bed with a temperature of 75°C. However, the micro void content and the presence of large and interconnected voids are lower in comparison to the 25°C and 50°C print bed samples.To investigate the influence of bed temperature on micro voids within the bead microstructure, an analysis was performed on the void volume fraction present in both the top and bottom beads of each 3D-printed two-layer composite. The measured results appear in Table 2, indicating that the void volume fraction exhibits variations within the microstructure of the two-layer beads printed at different bed temperatures and within each part's top and bottom beads. Furthermore, for a more comprehensive understanding of the micro void distribution within the microstructure of each bead, an evaluation of the void volume fraction was conducted along the coordinate direction of each bead. The results are presented in Figures (4-6).

Along the X-coordinate direction, the bottom bead consistently exhibits a higher void volume fraction than the top. Figure 4(b) presents the void volume fraction distribution within the microstructure of the two-layer bead printed on a bed temperature of 50°C. Like the bottom bead on the bed temperature of 25°C, a higher void volume fraction is observed along the edges of the bottom bead, which gradually decreases towards the bead center. In contrast, the top bead exhibits the opposite behavior, with a higher void volume fraction at the center and lower along the edges. Figure 4(c) illustrates the void volume fraction variation along the X-coordinate direction of the two-layer bead printed on a bed temperature of 75°C. Compared to the part printed on bed temperatures of 25°C and 50°C, a significant reduction in void volume fraction is observed in both the top and bottom beads.

Compared to the part printed on bed temperatures of 25°C and 50°C, a significant reduction in void volume fraction is observed in both the top and bottom beads. This observation indicates that increasing the bed temperature from 25°C to 75°C decreases the void volume fraction within the 3D printed part. Figure 4(d) demonstrates the void volume fraction along the X-coordinate direction of the part printed on a bed temperature of 100°C. It can be observed that the void volume fraction with a 100°C print bed is higher compared to the part printed on a bed temperature of 75°C but lower than those printed on bed temperatures of 25°C and 50°C. Additionally, within the top bead, the void volume fraction is higher near the center of the bead, while it is lower in the bottom bead. The findings in Figure 4230highlight the variation in void volume fraction within different regions of the composite part, depending on the bed temperature during printing.

The findings in Figure 4230highlight the variation in void volume fraction within different regions of the composite part, depending on the bed temperature during printing. Furthermore, the micro void distribution behavior within the top and bottom beads of each two-layer SCF/ABS composite exhibits inhomogeneous spatial patterns.Figure 4. Void volume fraction along the X-coordinate direction of two-layer SCF/ABS beads printed on beds with different temperatures: a) print bed temperature of 25°C, b) print bed temperature of 50°C, c) print bed temperature of 75°C, and d) print bed temperature of 100°C.Figure 5 presents the void volume fraction along the Y-coordinate direction (normal to the print surface) of the beads printed at different bed temperatures. Across all bottom beads, a lower void volume fraction is observed near the print bed, which gradually increases towards the center of the bead.

Across all bottom beads, a lower void volume fraction is observed near the print bed, which gradually increases towards the center of the bead. Notably, a significant increase in void volume fraction is detected just before the interface with the top bead, followed by a notable reduction at the interface itself. A comparative analysis of the void volume fraction within the microstructure of the top beads, printed at different bed temperatures, reveals distinct patterns. In all top beads, the void volume fraction is notably low, close to the interface between the top and bottom beads. However, the volume fraction gradually231increases, moving toward the center of the top beads. The void volume fraction is also consistently low near the surface across all additively manufactured parts.Figure 5.

The void volume fraction is also consistently low near the surface across all additively manufactured parts.Figure 5. Void volume fraction along the Y-coordinate direction of two-layer SCF/ABS beads printed on beds with different temperatures: a) print bed temperature of 25°C, b) print bed temperature of 50°C, c) print bed temperature of 75°C, and d) print bed temperature of 100°C.Figure 6 shows the volume fraction variation along the Z coordinate direction (corresponding to the 3D print direction) in two-layer SCF/ABS composites printed at different bed temperatures. Results indicate that unlike the variations observed along the X and Y-coordinate directions, the void volume fraction remains relatively constant along the print direction in both the top and bottom beads, as expected. Furthermore, this analysis reveals that the distribution of micro voids differs depending on the bed temperature.

Furthermore, this analysis reveals that the distribution of micro voids differs depending on the bed temperature. Specifically, in parts printed at bed temperatures of 25°C and 50°C, the void volume fraction is higher within the bottom beads. In contrast, the void volume fraction is higher within the top beads in parts produced at bed temperatures of 75°C and 100°C.232Figure 6. Void volume fraction along the Z-coordinate direction of two-layer SCF/ABS beads printed on beds with different temperatures: a) print bed temperature of 25°C, b) print bed temperature of 50°C, c) print bed temperature of 75°C, and d) print bed temperature of 100°C.The microstructure of each void within the two-layer SCF/ABS composites was subjected to further analysis to obtain detailed insights into their shapes.

Void sphericity, which is determined by comparing the surface area of the detected void with the surface area of a sphere with the same volume as the void, was employed to evaluate the shape characteristics of the micro voids. Note that a void with a perfectly spherical shape has a sphericity value of 1, while sphericity decreases as voids assume more irregular shapes. To ensure accurate results, voids of less than five voxels were excluded from the sphericity calculations, mitigating the impact of voxel size on the obtained sphericity values. A two-parameter Weibull distribution, given as:𝑓(𝑥) = (𝑏/𝑎)(𝑥/𝑎)(𝑏−1)(𝑒)−(𝑥 )𝑏 𝑎𝑥≥0 (1)was used to model the void sphericity distribution with parameters a and b representing the scale and shape parameters, respectively, which are presented in Table 3.

These figures reveal a considerable presence of voids with sphericity values below 0.5 (indicating irregular-shaped voids). In contrast, a smaller portion of voids exhibits sphericity values above 0.70 (indicative of more spherical-shaped voids). Furthermore, Figure 7(c) reveals that, among the four additively manufactured two-layer SCF/ABS parts, the sample printed on the bed with a temperature of 75°C exhibits a greater abundance of voids with higher sphericity values, indicating a more spherical shape. Specifically, a significant portion of the voids in this sample display sphericity values above 0.70. However, it is worth noting that voids within the sphericity range of 0.31-0.5 are still present in this particular sample. Figure 7(d) demonstrates that the void sphericity distribution within the composite printed on a bed temperature of 100°C resembles those printed on beds with 25°C and 50°C.

Figure 7(d) demonstrates that the void sphericity distribution within the composite printed on a bed temperature of 100°C resembles those printed on beds with 25°C and 50°C. The main distinction lies in the higher number of voids exhibiting a sphericity value greater than 0.7 in the sample produced at 100°C. µCT images also reveal distinctive microstructural characteristics at the interface between the top and bottom beads within the SCF/ABS-additively manufactured composites.Figure 8 provides 2D µCT images showing microstructural features at the interfaces of the printed parts.0.75 mm0.75 mm(a)(b)z0.75 mm0.75 mmx(c)(d)Figure 8.

2D µCT image of the microstructural features at the interface of two-layer SCF/ABS beads printed on beds with different temperatures: a) print bed temperature of 25°C, b) print bed temperature of 50°C, c) print bed temperature of 75°C, and d) print bed temperature of 100°C.235Figure 8(a) shows that at the interface of the part printed on a print bed with a temperature of 25°C, elongated voids, larger voids, and interconnected voids are observed alongside smaller voids. Figure 8(b) shows that the number of voids is reduced in the part printed on a bed with a temperature of 50°C. Although a few large voids and some interconnected voids are present, their occurrence is less frequent compared to the part printed at 25°C. No significant presence of large or interconnected voids is detected at the interface of the two-layer composite printed on the bed with a temperature of 75°C (Figure 8(c)).

No significant presence of large or interconnected voids is detected at the interface of the two-layer composite printed on the bed with a temperature of 75°C (Figure 8(c)). The voids observed in this case are considerably smaller and exhibit a more spherical shape than those in parts printed at 25°C and 50°C. Furthermore, Figure 8(d) highlights the presence of small voids, with only a few interconnected voids, at the interface of the two-layer composite 3D printed on a bed with a temperature of 100°C. This indicates that increasing the bed temperature from 25°C to 75°C causes a significant reduction in the average void volume fraction within the microstructure of the additively manufactured SCF/ABS composite and the number of interconnected voids at the interface.

It is worth mentioning that although the void sizes are larger at the interface of the part printed on a bed temperature of 100°C compared to the one printed on a bed with a temperature of 75°C, they are smaller with a significantly lower number of interconnected voids compared to those printed at bed temperatures of 25°C and 50°C.CONCLUSIONThis study aims to evaluate the influence of bed temperature on various aspects of the microstructure of two-layer SCF/ABS composites 3D printed with the Baylor LAAM system, employing µCT as the analysis technique. To achieve this, four composite parts were printed using identical printing process parameters but varying bed temperatures of 25°C, 50°C, 75°C, and 100°C. The void volume fraction within the microstructure of these additively manufactured composites was subsequently measured. The results show significant differences in the void volume fraction among the four printed parts.

The results show significant differences in the void volume fraction among the four printed parts. Furthermore, within each part, there was variation in void volume fraction between the top and bottom bead regions. Increasing the bed temperature from 25°C to 50°C reduced the void volume fraction from 12.47% to 12.06%, corresponding to a 5.33% decrease. Additionally, raising the bed temperature from 50°C to 75°C led to a substantial reduction in void volume fraction of 25.78% (from 12.06% to 8.95%). However, increasing the bed temperature from 75°C to 100°C caused an increase in void volume fraction from 8.95% to 11.11%, which warrants further investigation better to understand the underlying behavior and cause of this increase. This study also conducted a comprehensive analysis of void shapes by calculating their sphericity.

This study also conducted a comprehensive analysis of void shapes by calculating their sphericity. The findings revealed that increasing the bed temperature from 25°C to 75°C resulted in higher sphericity of voids, indicating a more spherical shape within the printed beads. This temperature increase was also associated with a reduced presence of interconnected voids. Conversely, as observed in the void volume fraction trends, raising the bed temperature from 75°C to 100°C decreased void sphericity. The microstructural characteristics, including void shape, were further examined at the interfaces of each printed composite part. The findings revealed that increasing the bed temperature from 25°C to 75°C substantially reduced the presence of interconnected voids. This reduction led to smaller void and a more spherical shape of the voids.

This reduction led to smaller void and a more spherical shape of the voids. At a bed temperature of 100°C, the voids were significantly smaller, and only a few interconnected voids were observed compared to those printed at 25°C and 50°C. However, it is worth noting that236the voids at the interface of the part printed at 100°C were larger than those at 75°C. These observations highlight the intricate relationship between bed temperature and void characteristics, indicating that higher temperatures contribute to forming smaller, more spherical voids with fewer interconnected voids.ACKNOWLEDGMENTThe authors would like to express their gratitude to the US National Science Foundation (grant #2055628) for their financial support and to Strangpresse for providing the Model 19 extruder used in this study.REFERENCES[1] T. Mulholl, S. Goris, J. Boxleitner, T.A. Osswald, and N. Rudolph, “Process-induced fiber orientation in fused filament fabrication”.

viewed PaperEFFECT OF THERMAL AGING IN THE FATIGUE LIFE OF HOT ISOSTATIC PRESSED AlSi10Mg FABRICATED BY LASER POWDER BED FUSIONS. T. Nabila,b, E. Arrietaa,b, R.B. Wickera,b, M. Benedictc and F. Medinaa,ba W.M. Keck Center for 3D Innovation, The University of Texas at El Paso, TX, 79968, USA; b Aerospace and Mechanical Engineering, The University of Texas at El Paso, El Paso, TX 79968, USA; cAir Force Research Laboratory, Wright-Patterson Air Force Base, Dayton, OH, 45433, USA;Corresponding author (snabil@miners.utep.edu) at The University of Texas at El PasoABSTRACTAlSi10Mg is a widely used material in the aerospace industry. Extended exposure to elevated temperatures can have a detrimental effect on it. In this work, multiple AlSi10Mg horizontal bars and vertical rods were fabricated using an L-PBF system. Following ASTM F3318-18, the material blanks were HIPed (Hot Isostatic Pressed). Emulating service temperatures, these blanks were aged at 177°C for 10, 100, and 1000h.

Following ASTM F3318-18, the material blanks were HIPed (Hot Isostatic Pressed). Emulating service temperatures, these blanks were aged at 177°C for 10, 100, and 1000h. Fatigue test specimens were machined down from the aged blanks. The machined specimens were subjected to a force-controlled fatigue test as per ASTM E466-15 with two stress levels: one within the elastic range (62MPa) of the material and another close to UTS (124MPa). The results indicated that even aging for 10h can dramatically reduce the fatigue life of the alloy.

The results indicated that even aging for 10h can dramatically reduce the fatigue life of the alloy. The work concludes with discussion on the reduction of fatigue life and visible progressive change in the ductility of the alloy with respect to the aging time.Keywords Laser Powder Bed Fusion(L-PBF), Fatigue Life, Thermal Aging, AlSi10Mg, Hot-Isostatic PressingNomenclatureL-PBF – Laser Powder Bed Fusion HIPed – Hot Isostatic Pressed YS – Yield Strength UTS – Ultimate Strength HT – Heat Treatment °C – Degree Celsius µm – Micrometer MPa- Megapascal Φ – Diameter HV- Hardness Value325INTRODUCTIONLightweight materials are always a primary focus in the modern aerospace industry. With the expansion of additive manufacturing technology, the possibility for the fabrication of complex and lightest weight structures has surpassed from rapid prototyping and moving forward to mass production[1], [2].

3D printing with powders has recently drawn the attention of engineers and scientists because of its versatile application in harsh environments[3]–[5]. Especially in the metal additive manufacturing sector- technologies like Powder Bed Fusion (PBF), Direct Energy Deposition (DED), Binder Jetting, etc., have constantly been used to develop complex geometries to aid modern aerospace and defense applications [6]–[9]. Thanks to improved powder materials and modern additive technologies, it is possible to manufacture lighter and cost-efficient components for the defense and aerospace sector.High-strength aluminum alloys- Due to their low weight, superior strength, and corrosive resistance properties, are widely used materials in the industry[11]. Especially AlSi10Mg, which has a density of ~2.66g/cm3 and a low melting point ranging from 570-590°C, is widely used to fabricate machined parts, heat exchangers, timing gears, automotive parts, and modern aerospace components[12], [13].

But, Al-Si parts are prone to porosities and defects, limiting the ability to obtain desired mechanical properties[14]–[19]. Nevertheless, it is possible to overcome such limitations by implementing powder bed fusion technology, especially Laser Powder Bed Fusion (L-PBF).Previous studies done with AlSi10Mg alloy with L-PBF have shown that it is possible to produce parts having nominal yield stress (YS) around 300MPa in contrast to almost 170 MPa seen for casted parts[20]. In addition, Numerous investigations of the post-heat treatment (HT) process of LPBF fabricated AlSi10Mg alloy, including stress relief (SR) anneals ranging around 200-540°C for 1 to 2 hours, solution treatments up to 550°C for 2 hours, Hot Isostatic Pressing (HIP) in 100MPa nitrogen environment at temperature up to 525°C for 2hours, and tempering or aging up to 530°C at holding times up to 10 hours along with various combinations were conducted.

These HT processes have resulted in a variety of mechanical properties in processed and post-processed L-PBF samples. Without HT, YS value was observed near 300MPa, and UTS value close to 480MPa while elongation was around 5 to 8%. After the HT, the YS transformed in a range of 80MPa to 170MPa with UTS in a range between 100MPa to 190MPa while varying the elongations ranging around 6 to 24%[21]–[31].This present study is aimed to investigate the fatigue performance of thermally aged AlSi10Mg alloy samples. These blanks were fabricated using LPBF and subsequently HIPed following ASTM F3318-18 at 510°C - 520°C, 100MPa for 3 hours. Then the HIPed blanks were thermally aged at 177°C for 0,10,100 and 1000 hours in the air.

Then the HIPed blanks were thermally aged at 177°C for 0,10,100 and 1000 hours in the air. Later, the specimens were machined for the fatigue test: one in the range of YS at 62MPa and another near UTS at 124MPa, and the fatigue endurance was evaluated.326METHODOLOGYPowder FeedAM 103C grade AlSi10Mg pre-alloyed powder (Inert gas atomized) from Valimet (Stockton, California) was used for the L-PBF system. The size and shape of the powder were analyzed using a Retsch Camsizer X2 (Haan, Germany) and shown in Figure 1. The size distribution ranged from 25 to 70 µm illustrated in where d10: 25.2µm, d50: 39µm, and d90: 59.2µm, respectively.

The size distribution ranged from 25 to 70 µm illustrated in where d10: 25.2µm, d50: 39µm, and d90: 59.2µm, respectively. Metallographic assessment of the internal section exhibited a micro-dendritic alpha-Al composition with secondary arm spacing of almost 1 micron, as shown in Figure 2.(a)(b)Figure 1: SEM images of Gas Atomized AlSi10Mg powder a) 100µm , b) 20µmFigure 2: Section View of Powder particle revealing micro dendritic structures327EOS M290 L-PBF systemsThe LPBF system used was EOS M290 (Kralling, Germany). It uses a 400W Yb-fiber laser and has an internal build volume of 250 x 250 x 325 mm[32]. A total of 48 (24 specimens in each orientation) Z-oriented vertical rods of 14mm x80 mm (Φ x H) and 24 XY-oriented bar blanks of 14 mm x 17mm x 76mm (L x W x H) were printed as test specimens.Print ParametersProcess parameters used for the LPBF system were pre-developed (as shown in Table 1) and able to produce builds with an average density of 2.65g/cm3.

This density was determined using a gas pycnometer which is equivalent to 99.25% of the density of 2.67g/cm3 EOS datasheet for AlSi10Mg printed in layers of 30µm with M290 systems[33].Table 1: Process Parameters used in EOS M290Laser Power Preheat Temperature of Bedplate Scanning Speed Hatch Distance Layer Thickness Print Pattern300W 80°C 1000 mm/s 130 µm 30 µm Striped with 7mm width and 20µm overlap distanceThermal Post-Processing: HIP and AgingThe As-build AlSi10Mg blanks and rods (a total of 48 specimens) were subjected to HIP as per ASTM F3318-18 standard[34].HIP was done at 515°C at 100MPa for 3 hours and cooled under an inert atmosphere to below 94°C. Followed by HIP, the blanks and rods were separated into four groups of twelve specimens: 6-Vertical (Z) and 6-Horizontal (XY), to be aged at 177°C for 0h, 10h, 100h, and 1000h.Fatigue TestThe specimens were machined from aged blanks for the fatigue endurance test following as per ASTM E466-15[35].

Based on previous findings, the YS and UTS for LPBF fabricated HIPed specimen was found to be 82MPa and 124MPa, respectively. The fatigue test was performed at two stress levels: one near the YS at 64MPa and the other at UTS at 124MPa. Three specimens followed the same orientation under each stress level. A sinusoidal waveform at 50Hz and R-value of 0.1 at room temperature of 22°C and 45% humidity was maintained throughout the fatigue test.328Hardness TestStruers Duramin-A300 (Struers, Cleveland, OH, USA) hardness tester was used, followed by Vickers (HV) scale for the measurements. 100gf and a dwell time of 5 seconds were applied for the test. Three numbers of indentations with separated nominal spacings of 1mm were done on two vertical (X and Y) and horizontal (Z) samples. The average of the indentations was used to calculate the hardness.

The average of the indentations was used to calculate the hardness. Satorius CP124S weight balance was used to determine the mass, and Accupyc II 1340 pycnometer was used to measure the volume by helium pycnometry. From these mass and volume measurements, density was calculated.Microstructure characterizationFor the characterization, the samples were derived from the unstressed portion of fatigue-tested specimens and cut into vertical and horizontal planes to analyze. Then ATM Opal 460 (Haan, Germany) mounting hot press was used to mount the cut samples in epoxy. Afterward, the samples were ground and polished with ATM SAPHIR 530 semi-automatic systems.

Then ATM Opal 460 (Haan, Germany) mounting hot press was used to mount the cut samples in epoxy. Afterward, the samples were ground and polished with ATM SAPHIR 530 semi-automatic systems. After that, the samples were submerged for 10 seconds in Keller’s etchant, and the microstructure was revealed using Olympus GX53 inverted optical microscope (Olympus Inc, Tokyo, Japan).RESULTS AND DISCUSSIONCorrelating Hardness value and Microstructure CharacterizationThe average density and hardness value of the specimens to printing orientation and aging hour is represented in Table 2. The density ranged from 2.65g/cm3 to 2.66g/cm3 for all test specimens, irrespective of the aging period. Hardness values measured from vickers micro indentation were consistent between horizontal and vertical planes, with lower HV values found for specimens aged 100h and 1000h. The microstructures of HIPed AlSi10Mg showed uniform distribution of Silicon (Si) particles in Figure 3.

The microstructures of HIPed AlSi10Mg showed uniform distribution of Silicon (Si) particles in Figure 3. No visible microstructural changes can be concluded after observing aging time from 0 hours to 1000 hours. However, specimens aged for 1000h showed finer Si particles of size below 3µm.Table 2: Density and hardness value based on printing orientationTest Matrix HIPed AlSi10Mg Density (g/cm3) Hardness (HV)0 hoursXYZ2.652.656059Thermally aged at 177°C10 hours ZXY100 hours ZXY2.652.662.652.66596049601000 hours ZXY2.652.654547329(a)(b)(c)(d)Figure 3: Microstructure of HIP AlSi10Mg showing uniformly displaying Si Particles.Aging time corresponds to a) 0, b)10, c)100 and d) 1000 hours at 177°CFatigue PropertiesThe evaluated fatigue properties for both: near YS at 62MPa and UTS at 124MPa for the aged HIP AlSi10Mg are represented in Table 3.

The specimens tested at 75% (62MPa) of the YS, were able to surpass five million cycles and, therefore, discontinued. There was no visible crack or failure indicators observed for the specimens tested, as shown in Figure 4.However, the fatigue specimens tested at UTS (124MPa) were prone to fracture at a different number of cycles in a visible correlation to their aging time, as shown in Table 3. The graph shown in Figure 5 presents the S-N plot for all the specimens tested at the UTS level (124MPa). The series of the graph consists of a group of three specimens each, and the names of the series indicate the HIP condition followed by aging time at 177°C (0h, 10h, 100h, and 1000h) and printing orientation (Z or XY). Progressive decay in the number of loading cycles was observed in Table 3 and Figure 5.

Progressive decay in the number of loading cycles was observed in Table 3 and Figure 5. The highest endurance was observed for the specimens that were not subjected to aging, and the330lowest was found for the 1000h aged specimens: with a maximum 99% reduction in the number of cycles with respect to unaged specimens. An important observation concluded from this observation is that the average number of cycles of unaged XY and Z specimens dropped from 1.31 and 1.23 million cycles to 0.8 and 0.87 million cycles, respectively, which represents that even aging for 10h can reduce the number of load cycle by 29% to 38% with respect to their printing direction as shown in Table 3. The specimens aged for 100h fractured in a range between 0.49 million to 0.73 million cycles in Figure 5(a) and Figure 5(b). On the other hand, the endurance of the specimens aged for 1000h was observed to range from 0.01 million to 0.02 million load cycles.

On the other hand, the endurance of the specimens aged for 1000h was observed to range from 0.01 million to 0.02 million load cycles. Surprisingly, one unaged Z specimen tested at UTS (124MPa) surpassed the threshold value of 5 million cycles and was discontinued (Table 3).

The hardness values derived for these specimens were 55.6, 58.7, and 54.7, respectively (Table 4). Therefore, no solid conclusion to differentiate can be yielded from the hardness values. Figure 6 represents the microstructure of the unaged Z-oriented specimens. Here, also no notably visible difference was observed in microstructural features.Table 4: Hardness Value of Z-oriented unaged SamplesZ- Orientation -0h aged 1 2 3 (Outlier)Hardness Value 55.6 58.7 54.7332(a)(b)(c)Figure 6: Microstructure of Z-oriented unaged Specimen (a) 1, (b) 2, (c) 3(outlier)Failure AnalysisThe fracture profiles shown in Figure 7 indicate that the specimens developed higher ductile features, such as reduced cross-section due to necking as well as cup and cone fracture modes with increased time of aging.

The 1000h aged specimen demonstrated the highest ductility in fracture profile: showing well-defined cup and cone fracture features.SEM fractography was further analyzed to observe the fracture surfaces of the Z and XY specimens at different aging times, as shown in Figure 8. It is observed that the fracture cross- section reduces with a longer aging time. Initially, it consists of a smooth texture with concentric lines merging to a crack initiation point and mapping out the directionality of crack propagation. Then, there is a rough surface which is the result of the rapid and final fracture. Higher aged specimens showed fewer features of these two textures due to ductility, which introduces more significant deformations accompanied by a more uniform stress distribution in the cross-section.

Therefore, a higher uniform texture was observed for the longer thermally aged specimens in Figure 8.333In addition, smooth and rough textures corresponding to crack propagation on both 0h and 100h XY specimens were less rough in Z specimens. This suggests although it is commonly believed that HIP post-processing removes microstructural anisotropy caused in LPBF systems, It cannot be entirely eliminated. The rougher surfaces in Z-specimens than XY-specimens verify it.High magnification SEM images were also examined where the rapid fracture occurred, as shown in Figure 9. It is seen that the fracture surface of the unaged Z specimen in Figure 9(a) had less voids and dimples in comparison to the surface of the Z specimen of 1000h aged in Figure 9(b).

It is seen that the fracture surface of the unaged Z specimen in Figure 9(a) had less voids and dimples in comparison to the surface of the Z specimen of 1000h aged in Figure 9(b). These differences were consistent and supported the previous observation on the fracture profiles and surfaces shown in Figure 7 and Figure 8, suggesting increased ductility for higher aged specimens.ZXYZXYZXYZ(a)(b)(c)(d)Figure 7: Fracture profiles at different aging time (Z and XY) (a) 0h, (b)10h, (c)100h, (d) 1000hZXY(a)XYZXYZXYZZXY(b)334XYXYZXYZXY(c)ZXY(d)Figure 8: SEM fractography at different aging time (Z and XY) (a) 0h, (b)10h, (c)100h, (d) 1000h(a)(b) (b)Figure 9: High Magnification SEM fractography at different aging time Z- orientation (a) 0h, (b) 1000h335CONCLUSIONThis research started with the goal of evaluating the fatigue performance of thermally aged L- PBF AlSi10Mg HIPed specimens.

Therefore, L-PBF AlSi10Mg samples (built-in Z and XY directions) were HIPed and thermally aged for 0, 10h, 100h, and 100h at 177°C emulating the service condition. After that, the specimens went through a fatigue test: one near YS at 62MPa and another in UTS at 124MPa for evaluation.Observing the microstructures, no evident change had been noticed at higher HIPed thermally aged samples. However, the fatigue test yielded dramatic results. For the fatigue test inspected near YS at 62MPa, the specimens in both directions lasted more than 5 million cycles and therefore discontinued.On the other hand, specimens tested at UTS at 124MPa showed a dramatic reduction in the number of cycles with respect to aging time. It has been found that even 10 hours of thermal aging can reduce the fatigue life of a specimen by 29 to 30%, which is a considerable drop.

It has been found that even 10 hours of thermal aging can reduce the fatigue life of a specimen by 29 to 30%, which is a considerable drop. Also, it was observed that 1000 hours of thermal aging reduce the fatigue life up to 98% of the fatigue life for the HIPed AlSi10Mg alloy fabricated from the LPBF machine.Diving into the fracture surface from the UTS fatigue specimens, it was seen that prolonged thermal aging gradually promotes ductility. Features like reduced cross-section increased necking, and cup and cone fracture profiles are progressive toward the 1000 hours of aging, validating the increased ductility of specimens due to aging.Furthermore, despite having identical fatigue endurance for both Z and XY specimens, the SEM fractography showed a noticeable difference between both. Z specimens had higher, less ductile textures in SEM fractography than the XY specimens. The ductility was less prominent for the 0h and 10h than the 100h and 1000h specimens.

Z specimens had higher, less ductile textures in SEM fractography than the XY specimens. The ductility was less prominent for the 0h and 10h than the 100h and 1000h specimens. This suggests a future investigation of whether hot isostatic pressing can entirely eliminate the microstructural features of the inherent printing and to what degree of homogenization is allowed as ideal conditioning.ACKNOWLEDGMENTThis research presented here was conducted at The University of Texas at El Paso within the W.M. Keck Center for 3D Innovation (Keck Center), a 13,000-sq. ft. state-of-the-art additive manufacturing facility. Additionally, equipment housed within the Department of Metallurgical, Materials, and Biomedical Engineering was utilized. Support for the research was provided, in part, by Mr. and Mrs. MacIntosh Murchison Chair I in Engineering with strategic investment in the development of this work by the UTEP Aerospace and Mechanical Engineering Department and the W.M.

This study compares the microstructure and tensile properties of two laser powder bed fused Al alloys, namely AlSi10Mg and Scalmalloy, in different heat-treated (HT) conditions i.e., stress-relief, T6, and hot isostatic pressing (HIP), as well as the non-heat- treated condition. The microstructures were examined using scanning electron microscope, and the mechanical properties were evaluated using uniaxial tensile testing. For AlSi10Mg, Si- networks were observed to break down, and Si- and Fe-rich particles precipitates form during HIP followed by T6. For Scalmalloy, the density of nanometer-sized intergranular Al3(ScxZr1-x) precipitates increased after only HIP.

For Scalmalloy, the density of nanometer-sized intergranular Al3(ScxZr1-x) precipitates increased after only HIP. Furthermore, it was observed that the addition of HIP improved the tensile strengths of both alloys as compared to their AM as well as wrought counterparts in similar HT conditions.Keywords: Laser powder bed fusion (L-PBF), AlSi10Mg, Scalmalloy, Hot isostatic pressing (HIP), Microstructure, Tensile propertiesIntroductionAdditive manufacturing (AM) allows the fabrication of complex shapes by layer-by-layer process [1]. The laser powder bed fusion (L-PBF) is a widely used AM technology for various Al alloys such as AlSi10Mg. Al alloys are well known in different industries, including automotive, aerospace, construction, and packaging, due to their superior material properties such as light- weight, good corrosion resistance, high thermal conductivity, excellent machinability, etc. [2–4].

[2–4]. AM of Al alloys could be more challenging than other metals because Al powder particles are naturally light and have high moisture absorption, high thermal conductivity, and high oxidation341tendency, all of which may affect the fabrication process and therefore, resulting in detrimental process-induced defects on the mechanical properties of the AM Al components [5].AlSi10Mg is one of the most popular and widely used Al alloys in AM industry. Characterized by low coefficient of thermal expansion (CTE), AlSi10Mg components can be manufactured by AM with great dimensional precision and low thermal stresses. Post-processing heat treatments (HT) are frequently used for AM metals to partly alleviate the detrimental effect of some characteristics.

Post-processing heat treatments (HT) are frequently used for AM metals to partly alleviate the detrimental effect of some characteristics. These include the dendritic microstructure that improves the mechanical properties of alloys [6].There are several research in the literature studying the effect of different thermal post- processes on microstructure and mechanical properties of L-PBF AlSi10Mg [7–13]. The conventional HT for the AM AlSi10Mg alloy consists of two steps: (1) T4: solid solution (at 540°C for 6 hr, then water quench) followed by (2) T6: artificial ageing (at 160°C for 4 hr, then air cool) [7–9]. The T4 HT could have complete homogenization effect on microstructure while T6 could improve the strength of the L-PBF AlSi10Mg alloy. Nonetheless, EOS merely recommends a low temperature stress relieving (SR) at 300°C for 2 hr [12]. It should be noted that the hot isostatic pressing (HIP), as common HT for certain AM metals, could also be used for L-PBF AlSi10Mg.

It should be noted that the hot isostatic pressing (HIP), as common HT for certain AM metals, could also be used for L-PBF AlSi10Mg. But having HIP along may cause a significant strength reduction with remarkably increased ductility for L-PBF AlSi10Mg [13].Scalmalloy is a Sc- and Zr- modified high strength Al alloy, which has received a lot of attention by AM industries [14–18]. The L-PBF Scalmalloy have shown a bimodal grain structure consisting of nano-size equiaxed grains as well as fine columnar grains [15]. The conventional annealing HT for Scalmalloy is a low temperature HT at 325°C for 4 hr. [17, 18], resulting in nanometer-sized Al3(Sc, Zr) precipitation along the grain boundaries which prevents grain growth [16].HIP is widely recognized for its ability to reduce volumetric defects in the material or even removing them. HIP alone or followed by appropriate HT schedule could modify the tensile properties and which in turn, improves the fatigue resistance of the materials.

HIP alone or followed by appropriate HT schedule could modify the tensile properties and which in turn, improves the fatigue resistance of the materials. In this article, the effect of HIP treatment for both L-PBF AlSi10Mg and L-PBF Scalmalloy were investigated and compared with the typical conventional HT on microstructure and tensile properties of the alloys.Experimental ProcedureMaterial and FabricationCylindrical bars (15 mm in diameter by 90 mm in height) of AlSi10Mg and Scalmalloy were fabricated using L-PBF techniques, by Valimet [19] and Heraus Additive Manufacturing [20], respectively. The chemical compositions of AlSi10Mg and Scalmalloy powders are presented in Table 1. Furthermore, Table 2 presents the process parameters used in this study for the fabrication of cylindrical bars, provided by the manufacturer. Small coupons (5 mm thickness) were cut from cylindrical bars for further microstructural analysis.342Table 1.

Small coupons (5 mm thickness) were cut from cylindrical bars for further microstructural analysis.342Table 1. Chemical composition of AlSi10Mg powder, provided by Valimet [19] and Scalmalloy powder, provided by Heraeus Additive Manufacturing [20].Si Mg FeCu MnTiZnNi OthersAlAlSi10Mgwt. %10.2 0.3 0.09 0.002 0.002 0.011 0.004 0.006 < 0.10BalScalmalloyMg Sc Zr Mn wt. %SiFeZn4.29 0.7 0.32 0.59 0.053 0.099 0.004 0.002O AlCuTi0.018V0.01wt. %0.05 BalTable 2. Process parameters used for fabricating the L-PBF AlSi10Mg and L-PBF Scalmalloy.MaterialLaser power (W)Scan speed (mm/s)Hatch distance (µm)Layer thickness (µm)AlSi10Mg370148012060Scalmalloy380142514530The multi-steps heat treatment schedules used in this study for the both alloys are listed in Table 3, which are also schematically shown in Figure 1.

Following the ASTM F3301-18a [21] and AMS 2771F [22] standards, the heat treatment schedules were selected. The heat treatment was performed under an argon-atmosphere in an electric furnace.Table 3. Different post-process HTs used in this study.MaterialAlSi10MgHTSR+HIP+T6SR+T6SR270°C/1.5 hr/FC270°C/1.5 hr/FCHIP520°C/100 MPa/2 hr/FC---T6 (Solution + Aging) 540°C/6 hr./WQ + 160°C/4 hr/AC 540°C/6 hr./WQ + 160°C/4 hr/ACSR325°C/4 hr/FC------ScalmalloyHIP---325°C/100 MPa/4 hr/FC---Notes: FC: Furnace Cool, WQ: Water Quench, AC: Air Cool343Figure 1. Schematic diagrams of HT schedules conducted on (a) L-PBF AlSi10Mg and (b) L- PBF Scalmalloy.Microstructure characterizationMicrostructural characterization was conducted in the transverse direction (TD) plane, i.e., parallel to the build direction.

The metallography procedures were performed according to the ASTM-E3 [23] including cold-mounting, grinding with SiC sandpapers with grits sizes ranging from 320 to 4000, and polishing to a mirror-finished surface using polishing cloths and colloidal silica suspensions size of 0.05 µm. Electron backscatter diffraction (EBSD) analysis and electron channeling contrast imaging (ECCI) were conducted for microstructure characterization using a Zeiss Cross Beam 550 scanning electron microscope (SEM).Mechanical testingThe cylindrical bars were machined upon HTs to the final geometry of uniaxial tensile tests (see Figure 2) according to the ASTM E8 [24]. The room temperature tensile tests were performed using an MTS landmark servo-hydraulic load frame with a load cell of 100 kN. The quasi-static tensile tests were conducted in two steps: strain-controlled followed by displacement-controlled step.

The quasi-static tensile tests were conducted in two steps: strain-controlled followed by displacement-controlled step. An MTS mechanical extensometer was attached to the specimen in the gage section to record the strain. However, due to the limited travel range of the extensometer, it was removed at ~0.035 mm/mm strain by pausing the test, after which the test was resumed in the displacement-controlled344mode up to fracture. At least three tests for each HT condition were conducted to ensure the repeatability of the results.Figure 2. The geometry of the tensile specimens following ASTM E8 [24]Results and DiscussionMicrostructure analysisThe Inverse Pole Figure (IPF) maps for the non-heat treated (NHT) and HT of the both L- PBF AlSi10Mg and Scalmalloy obtained from EBSD analysis are presented in Figure 3 (a) to (c) and (d) to (f), respectively. The IPF maps were obtained in the transverse direction (TD) plane (i.e., parallel to the build direction.

The IPF maps were obtained in the transverse direction (TD) plane (i.e., parallel to the build direction. On one hand, for L-PBF AlSi10Mg, the grain structure was mostly columnar which is typical for AM metal alloys [25]. As shown in Figure 3(a)-(c), the average grain sizes were affected by the HTs, i.e., ~ 2.9 μm for the NHT condition, ~ 5.3 μm for SR+T6, and for ~ 10.6 μm SR+HIP+T6.On the other hand, the L-PBF Scalmalloy showed much finer grain structure with bi-modal morphology consisting of (1) nano-sized equiaxed fine grains (FG) along melt pool boundaries and (2) coarser grain (CG) away from melt pool (see Figure 3(d)-(f)). The average grain sizes of the L-PBF Scalmalloy in different HT conditions did not show significant changes, ranging from 3.8, 2.8 and 5.6 μm for NHT, SR, and HIP conditions, respectively. The presence of such fine grain structure in NHT condition could be attributed to the rapid solidification rate during the L- PBF process [26].

The presence of such fine grain structure in NHT condition could be attributed to the rapid solidification rate during the L- PBF process [26]. Upon HTs, similar fine grain structure is maintained which is attributed to the precipitation of nano-size Al3Sc and Al3Zr particles, which can act as grain growth inhibitors [27]. It is worth noting that the actual value of the average grain sizes for the L-PBF Scalmalloy may be significantly smaller than what was reported using EBSD analysis due to step size limit (0.43 μm).The BSE micrographs obtained by ECCI technique for both the alloys are presented in Figure 4 in all the different HT conditions investigated in this study. As shown in Figure 4(a), NHT condition of L-PBF AlSi10Mg consisted of the cellular Si network. Upon HTs, these networks break down and form bulky Si-rich and needle shaped Fe-rich particles at grain boundaries (GB) and grain interiors (see Figure 4(b)-(d)) [7,28].

Upon HTs, these networks break down and form bulky Si-rich and needle shaped Fe-rich particles at grain boundaries (GB) and grain interiors (see Figure 4(b)-(d)) [7,28]. Furthermore, due homogenization of high temperature solutionizing (T4) or HIP stages, the micro-segregation of Si and Mg elements that were observed in the NHT specimens (see Figure 4(a)) are completely dissolved back into the α-Al matrix, resulting in supersaturated solid solution (SSSS). It should be345noted that during the subsequent age hardening T6 step, the nanometer-sized Mg2Si particles could precipitate from the SSSS matrix, which are only detectable by transmission electron microscopy (TEM) therefore are not shown over here [29].Figure 3. Inverse pole figure (IPF) maps obtained by EBSD analysis in the TD plane in (a) to (c) for the L-PBF AlSi10Mg and in (d) to (f) for the L-PBF Scalmalloy in different heat treatment conditions investigated in this study.346Figure 4.

BSE micrographs obtained in TD for (a)-(d) L-PBF AlSi10Mg and (e)-(g) L-PBF Scalmalloy in different heat treatment conditions.For L-PBF Scalmalloy, upon SR and HIP, the fine inter dendritic regions are partially dissolved back into the α-Al matrix (see Figure 4(e)-(f)). However, as discussed earlier, the formation of small Al3(Sc, Zr) phases at GBs during the low temperature SR and HIP as GB boundary strengthening particles help the alloy maintain the prior fine grain structure observed in NHT condition.Tensile propertiesThe uniaxial tensile properties including yield strength (YS), ultimate tensile strength (UTS) and percent elongation to failure (%EL) are presented as a column chart in Figure 5 (a) and (b) for L-PBF AlSi10Mg alloy and L-PBF Scalmalloy, respectively.

In addition, the tensile properties of cast AlSi10Mg [30] and wrought Scalmalloy [31] are also included for comparison purposes.It can be seen in Figure 5 (a) that tensile strengths (e.g., YS and UTS) and %EL of L-PBF AlSi10Mg in SR condition was significantly higher than those of cast counterpart by ~87%, 44% and 170%, respectively, which could be attributed to the typical columnar fine grain structure observed for the former (see Figure 4(b)) [32-33]. As discussed earlier, upon T6, due to formation347of Si-rich and Fe-rich particles precipitate at GBs and grain interiors as well as small Mg2Si precipitates [7, 28], the tensile YS of the alloy is increased by ~12% which is compromised with a decrease in EL [10, 34].

Moreover, as it can be seen the intermediate HIP prior to T6 heat treatment showed relatively small impacts on tensile strengths of the alloy while it resulted in significant improvement on %EL of the alloy which could be attributed to the effect of HIP on closing the process induced pores [13,35].It can also be observed that tensile strength of L-PBF Scalmalloy in SR condition is significantly higher as compared to the wrought counterpart (see Figure 5(b)). The YS and UTS are increased by ~60% and ~31%, respectively, which could be attributed to the nano-sized grain structure as well as Al3(Sc, Zr) precipitates [33,36]. However, the EL was compromised and significantly decreased from 20% to 1.4%. Upon further HT, the tensile properties were affected insignificantly.348Figure 5.

However, the EL was compromised and significantly decreased from 20% to 1.4%. Upon further HT, the tensile properties were affected insignificantly.348Figure 5. Bar chart representation of the tensile properties of (a) AlSi10Mg and (b) Scalmalloy under different heat treatment conditions investigated in this study.ConclusionsIn this study, the effect of different heat treatments on microstructure and tensile properties of AlSi10Mg and Scalmalloy fabricated by L-PBF are investigated. The test specimens underwent different post thermal processing including stress-relief (SR), followed by hot isostatic pressing (HIP) and T6 for L-PBF AlSi10Mg. For L-PBF Scalmalloy, the heat treatments were SR and HIP. The microstructure was characterized between each two steps of the multi-step heat treatments. Furthermore, the tensile properties were evaluated through the quasi-static tensile testing. The experimental observations from this study can be concluded as follows:1.

Furthermore, the tensile properties were evaluated through the quasi-static tensile testing. The experimental observations from this study can be concluded as follows:1. Grain growth was seen for both L-PBF AlSi10Mg after the different post process heat treatment. For L-PBF AlSi10Mg, the grains are typically columnar and after SR+HIP+T6, the grain size increased even further than the SR+T6 condition. A bi-modal grain structure is observed consisting of nano-sized equiaxed grains along the melt pool boundaries for L- PBF Scalmalloy, which become more finer with partially elongated grains after HIP process.2. A cellular network of Si phase was seen for non-heat-treated (NHT) L-PBF AlSi10Mg. The Si-networks dissolved after SR+T6 heat treatment and even after SR+HIP+T6, Si- particles and Fe-rich particles precipitates.

The Si-networks dissolved after SR+T6 heat treatment and even after SR+HIP+T6, Si- particles and Fe-rich particles precipitates. For L-PBF Scalmalloy, the Sc and Zr elements349are accumulated and distributed in the matrix after SR, and even after HIP process the strong Al3(Sc, Zr) phase precipitates.3. The tensile strength and ductility increased after SR+HIP+T6 for L-PBF AlSi10Mg compared to SR+T6 condition because of fine Si- and Fe-rich particles precipitates. The L-PBF Scalmalloy showed higher strength and ductility after HIP.AcknowledgmentThis research is partially supported by the National Aeronautics and Space Administration NASA under Cooperative Agreement No. 80MSFC19C0010. This paper describes objective technical results and analysis.

fect of Volume of Heat Sink on Process and Physical Properties of Parts Built by Welding Based SFFZ. Jandric, J. H. Ouyang and R. KovacevicResearch Center for Advanced Manufacturing Department of Mechanical Engineering, Southern Methodist University 1500 International Pkwy. Ste. 100, Richardson, TX 75081AbstractA new numerical simulation of the effect of the volume of the heat sink on the welding–based deposition process is performed. For this purpose, the ANSYS parametric design language (APDL) is applied. Due to the complex internal and/or external shapes of the designed three- dimensional (3D) part, different heat transfer conditions are met during the building process. The influences of the different heat transfer conditions on the physical part properties are also investigated. The influence of the volume of the heat sink on the process and on the physical properties is significant and can not be neglected.

The influence of the volume of the heat sink on the process and on the physical properties is significant and can not be neglected. Extensive experiments are designed and executed in order to verify the conclusions derived from the finite elements model results and to investigate the material properties of the built part.IntroductionThe majority of the solid freeform fabrication (SFF) techniques use non-metallic materials for building parts. SFF based on welding is technique capable of making the 3D metallic parts. Any weldable metal can be used as a material for deposition. Also, several research-teams [1 – 9] have shown that this technique is capable of making diverse part shapes. Spencer et al. made a vertical wall and a hollow box. Ribeiro et al. made a ‘chimney’ shape, a ‘bow tie’ shape, and a ‘pint glass’ shape. Kovacevic et al.

Spencer et al. made a vertical wall and a hollow box. Ribeiro et al. made a ‘chimney’ shape, a ‘bow tie’ shape, and a ‘pint glass’ shape. Kovacevic et al. also made a thin vertical wall, cylinders, cubical 3D parts with straight channels, a 3D part with complex geometry and a 3D network of the cooling/heating conformal channels.The biggest challenge for this process is to control the process parameters. Namely, during the deposition process, it was apparent that process parameters cannot remain constant since the deposited layers were not smooth. In some instances, the beads were flat and wide, but in others they were tall, and looked like the heat input was insufficient. We looked for possible reasons for such huge fluctuations, and the influence of the geometry of the previously deposited layers was the only answer. This phenomenon is further investigated by modeling it with finite element method (FEM).

This phenomenon is further investigated by modeling it with finite element method (FEM). For this purpose, the ANSYS parametric design language (APDL) is applied.A substantial amount of research has been documented for modeling and understanding the welding process. Rosenthal [10] and Rykalin [11] developed a solution to the heat equation for point and line heat sources on idealized geometries such as infinite plates. Because of computational difficulties encountered when attempting to include inelastic strains in an analysis in which rapid temperature changes take place, some analytical work was carried out that goes beyond Tall’s one-dimensional analysis of longitudinal stress [12].331A determination of the complete thermo-mechanical response due to welding requires a full three-dimensional inelastic analysis of mechanical behavior, accompanied by a computation of the three-dimensional transient temperature distribution.

In the last decade, models for nonlinear transient finite element heat transfer analyses of the welding process have been developed [13 - 15]. Numerical modeling requires, at the beginning, considerable time and effort, but once the model is obtained, it is easy to change any welding parameter, welding conditions, type of the material, etc. The objective of this research is to numerically simulate the welding process during the deposition process and to evaluate the physical characteristics of the deposited material as a function of heat sink volume.Finite Element ModelIn this study, the commercial finite element software (ANSYS) was employed. It provides a convenient means of numerically modeling the welding process. The model is discretized into individual finite elements that collectively approximate the shape of the model. A 3D thermal solid element type with a 3D thermal conduction capability is used.

The model is discretized into individual finite elements that collectively approximate the shape of the model. A 3D thermal solid element type with a 3D thermal conduction capability is used. The element has eight nodes with a single degree of freedom, temperature, at each node. The element is applicable to a three- dimensional, steady-state or transient thermal analysis. In our model, we mostly used a ‘mapped’ mesh so that the mesh has a regular pattern with obvious rows of elements (Fig. 1). Also, we used transition-mapped hexahedral meshing, by specifying line divisions on the side edges of the volume such that the divisions permit a transition-mapped hexahedral mesh, i.e.

1). Also, we used transition-mapped hexahedral meshing, by specifying line divisions on the side edges of the volume such that the divisions permit a transition-mapped hexahedral mesh, i.e. the elements at the top layer of the substrate are defined with a significantly smaller height than the elements at the very bottom of the substrate, since the temperature gradients in the vicinity of the top surface, where a heat flux is applied, is much higher.For this model, the thermal fields and thermal history is required, so a transient thermal analysis must be performed. This requires an integration of the heat conduction equation with respect to time.

This requires an integration of the heat conduction equation with respect to time. The basic FEM equation for the thermal problem can be derived from the thermal equilibrium equation: (cid:6) ( ) { }  T C T+ ( ) { } { } { } + Q =K TTV(1)where ] [ ]T C [ ]N = the shape function matrix [ [ ][ ]T = ∫ k B B dV K (cid:6) { }T and { }T temperature rate vector, respectively, and { }Q = the heat flux vector][ [ c N N dVρ= ∫is the heat capacity matrixV]is heat conduction matrixTVare nodal temperature vector and nodalFig. 1 Meshed model.Modeling ResultsThe isotherms obtained with the 3D thermal analyses for different time intervals are given in Fig. 3a – 3f. Points of interest are located, where the surrounding mass of the material varies between33212% and 100% of the maximum possible (points are marked with the letters A – F, see Fig. 2). The heat input is kept constant and moves across the top surface at a steady rate.

2). The heat input is kept constant and moves across the top surface at a steady rate. In the model, heat flux travels at speed of 23 cm/min. That is the actual speed that was used during depositon of material in welding-based SFF.From these figures, it can be observed the change of temperature with different heat transfer conditions. Consequently, the size of the molten pool and the size of the heat affected zone alter as well. The central temperature contour in each image represents a critical temperature of 1530°C the melting temperature. Hence, this contour denotes the boundary of the molten pool. In Fig. 4 is shown a 45° corner together with the molten pool and the heat affected zone.asFig. 2 Geometry of the model.a)b)c)f) Fig. 3 Temperature fields as a function of volume of heat sink. In Fig. 4a the heat input is far from the corner (45°), so that the volume of heat sink is half of the maximum possible one.

3 Temperature fields as a function of volume of heat sink. In Fig. 4a the heat input is far from the corner (45°), so that the volume of heat sink is half of the maximum possible one. At this instant, there is barely enough heat to melt material and consequently a small molten pool is formed. In Fig. 4b, the heat input is closer to the corner, the volume of the heat sink is decreased, and, the amount of the melted material is enlarged. In the Fig. 4c, the heat input is placed above the corner itself. The volume of the heat sink is four times smaller than in the first case, and as a result, the molten pool and heat affected zone are much larger.d)e)a)b)c)Fig. 4 Influence of the geometry on the size of the molten pool and the heat affected zone.333Although the difference between these three cases is obvious, it is not easy to quantify it through the numerical values. For this purpose, the temperature history was used.

For this purpose, the temperature history was used. Temperature history from the model is verified experimentally by employing the thermocouples (see Fig. 5). For each thermocouple a hole is drilled in the appropriate edge. Each hole is drilled vertically from the bottom toward the top of the sample. The depth of the each channel is for 1.5 mm less than the height of the sample. To isolate thermocouples from noise that originates from the arc, in the bottom of each channel, thermally conductive and electrically insulating high temperature cement is placed. The thermocouples and thermally conductive cement are inserted together,Fig. 5 Embedded thermocouples.Fig. 6 Temperature history.ModelModelExperimentExperimenta) 45° anglec) 90° angleFig. 7 Experimental vs. modeling temperature history.and left for the cement to cure.

6 Temperature history.ModelModelExperimentExperimenta) 45° anglec) 90° angleFig. 7 Experimental vs. modeling temperature history.and left for the cement to cure. In order to compare the temperature history from the model with experimental results, nodes 1.5 mm beneath the top surface, from each corner, are selected. The temperatures from the model, for all angles, are recorded in time (Fig. 6). It could be seen that smaller heat sinks achieve a much higher temperature, and accordingly the molten pool enlarges. These results are also in good standing with the results obtained by Kamala et al. [16], where they also showed that the temperature increases when the heat flux is approaching a smaller heat sink. Direct comparisons of experimental and modeling results, for angles of 45° and 90° are shown in Fig 7. a) and b), respectively. It is evident that the experimental results are lower than those from the model.

a) and b), respectively. It is evident that the experimental results are lower than those from the model. There could be several reasons for this difference: the model has several assumptions (such as the heat input remains constant during the process while in reality fluctuations in the current and voltage cannot be avoided, so, to minimize the error, the mean value is used; the workpiece is at a uniform temperature before the deposition starts; the body is exposed to a uniform forced convection; the surface-to-surface contact between the workpiece and ‘holding plate’ beneath is ideal, so that heat is transferred by conduction throughout the whole contact area; the effect of gas diffusion in the weld pool was not considered), as well as a334difference in the conductivity of the cement and mild steel, the fact that readings from thermocouples are too slow to capture the temperature peak, etc.

In general, we can conclude that results of the model are in good standing with the experimental results. This proves that the model is set properly and that can rely on its results. This difference in the temperature leads to the well known fact: if the same material is heat treated in different way, a different microstructure is to be expected.Influence of the different volume of the heat sink on the material propertiesTo investigate the influence of the volume of the heat sink on the material properties of the 3D part built by welding-based SFF, several experiments are designed and conducted. Material is deposited on the 3D pre-machined parts, which are of the same shape and size as the finite element model (shown previously). This shape simulates different heat transfer conditions by altering the size of the heat sink. Thus, the size varies from 100% for the bead placed at the flat top surface to the 12.5% for the bead placed at the 45° corner.

Thus, the size varies from 100% for the bead placed at the flat top surface to the 12.5% for the bead placed at the 45° corner. During each experiment, the process parameters are kept constant. Overall, nine experiments are performed, while current is changed from 80 A to 160 A, with increments of 40 A, and wire feeding speed is changed from 40 cm/min to 120 cm/min, with increments of 40 cm/min. Mild steel welding wire with a diameter of 0.9 mm is used.]m m[b B,p D,w B8642004590135 180 225 270 315 360AngleFig. 8 Geometry of the bead vs. angle.B wD pB bAfter depositing beads, they are cross- cut along the lines A-A, B-B, C-C, D- D, E-E and F-F (see Fig. 2) in order to obtain information about the weld bead geometry. After polishing and etching the cross-sections, all beads are measured and the averaged values for bead width, bW, bead build-up, bB, and depth of penetration, dP, are shown in Fig. 8.

After polishing and etching the cross-sections, all beads are measured and the averaged values for bead width, bW, bead build-up, bB, and depth of penetration, dP, are shown in Fig. 8. While bead width and depth of penetration decreases with increasing the volume of the heat sink (from 45° to 360°) 2.15 and 2.06 times, respec- tively, bead build-up remains the same.a) 45°b) 270°c) 360°Fig. 9 Cross-sections of the bead at different angles.Images of the whole bead cross-sections at 45°, 270° and 360° are shown in Fig. 9. Enlarged images of the beads are shown in Fig 10. We chose to show only the two most extreme cases (45°- Fig. 10a and 360°- Fig 10 c ) and one at the angle of 270° (Fig. 10 b). Images are scaled equally, and they can be compared directly. Images of the cross sections are taken in the middle335part of the beads, as it is indicated with vertical lines in Figs. 9 a – c.

Images are scaled equally, and they can be compared directly. Images of the cross sections are taken in the middle335part of the beads, as it is indicated with vertical lines in Figs. 9 a – c. The maximum size of the bead width, bead build-up, depth of penetration and heat affected zone are obtained.Each bead consists of the melted zone (marked as I in Figs. 10) and the heat affected zone (marked as II in Figs. 10). The melted zone of the bead (the upper part) consists of the build-up and depth of penetration parts. From Fig. 10 a, b and c it is obvious that the overall lengths of the beads are different. The longest bead is the one at 45°, while the shortest is at 360°. Logically, the longest bead has the longest melted as well as heat affected zones. This is a direct consequence of the highest temperature that this part is exposed to.

Logically, the longest bead has the longest melted as well as heat affected zones. This is a direct consequence of the highest temperature that this part is exposed to. Since the heat transfer conditions are much more different between 45° and 270°, than between 270° and 360°, beads at 270° and 360° are more alike.Also, from these images we can observe the size and the structure of grains in each zone. Again, the bead at 45° is unique in that it has the coarsest grains in the melted zone. Its grains are very coarse columnar grains, which tell us that heat was conducted mostly in the vertical direction. The growth direction during solidification is approximately perpendicular to the solid/liquid interface and is aligned to the heat-flow direction. Consequently, the overall is microstructure of controlled by the interplay between the thermal field the growth crystallography.

Consequently, the overall is microstructure of controlled by the interplay between the thermal field the growth crystallography. Besides the epitaxial growth of the original grain boundaries there exist a large number of fine columnar subgrain boundaries that are also rich in the precipitates. At 270°, columnar grains are still dominant, but their length is visibly smaller and they are wider. Since at this angle more material exists around the bead, grain growth occurs not only against vertical heat flow. At 360°, heat is conducted in all directions and as a result very fine columnar grains are formed. The change in the solidification morphology could have been due to the differences in thermal and fluid conditions in the filler-deposited molten pool. It should be noted that all three deposited beads exhibit equiaxed grains at the top part of the bead. The influence on the geometry is clear.

It should be noted that all three deposited beads exhibit equiaxed grains at the top part of the bead. The influence on the geometry is clear. Not only is it true that, with the smaller heat sink, the temperature rises and more material is melted resulting in a larger heat affected zone, but it is also true that the material properties are different as well. This means that the physical properties will not be uniform if the process parameters are constant regardless of the geometry.IIIIIthe solidified zoneIIinthe melt pool andb) 270°c) 360°II= 0.76mma) 45°Fig. 10 Enlarged beads at different angles.Fig. 11 Original material.336a)a)a)An image of the original material is given in Fig. 11. In addition, each of these beads is observed with the larger magnification. Results can be seen in Fig. 12 (45°), Fig. 13 (270°) and Fig.

11. In addition, each of these beads is observed with the larger magnification. Results can be seen in Fig. 12 (45°), Fig. 13 (270°) and Fig. 14 (360°).b)c)d)b)c)d)b)c)d)All three beads consist of the build-up zone, the penetration zone and the heat affected zone. Although the grain-size shown in the figures is much different, the phase constituents at different angles are similar. The heat affected zone exhibits fine martensite and α-ferrite. The is composed of coarse penetration zone columnar dendrites. Some fine cementite precipitate at the grain boundaries along a certain crystal plane as shown in Figs. 12-14 (b). The segregation of carbon atoms at the grain boundaries leads to the precipitation of the carbides, which may distinctly affect mechanical properties of the beads.≡ 84.7µme) Fig. 12 45° Fig. 13 270° Fig. 14 360°New Experimental Set-upAll of the aforementioned experiments are new experimental set-up (shown in Fig.

12 45° Fig. 13 270° Fig. 14 360°New Experimental Set-upAll of the aforementioned experiments are new experimental set-up (shown in Fig. 15) where a 6-axis robot is added. A new torch and a wire feeding mechanism are designed the robot’s specifications. This set-up allows us ‘more freedom’ in parts design, since robot can handle more advanced geometry. Thus, we built a turbine blade (see Fig. 16e). In Fig. 16 we can see the procedure of making the part – from designing the part in CAD software then generating the cross-section (Fig. 16b) and the torch path (Fig. 16c) with the new software “Success”, which is also recently developed by our research team, and at the end, depositing material (Fig. 16d) and obtaining the final accuracy by milling the top and side surfaces (Fig. 16e).performedonourto accommodate(Fig. 16 a),Fig.

16d) and obtaining the final accuracy by milling the top and side surfaces (Fig. 16e).performedonourto accommodate(Fig. 16 a),Fig. 15 Experimental Set-up.337a) 3D CAD modelb) Cross-sectionc) Path planningd) Turbine blade before millinge) Turbine blade after millingFig. 16 Procedure for making turbine blade by SFF based on welding.ConclusionsA three-dimensional finite element model of the transient, temperature-dependent heat transfer problem has been developed. This model affords a better understanding of the fluctuations that appear when a constant heat input is applied during the deposition process. The temperature fields are shown for the ‘critical’ angles. For these angles temperature history from the model is recorded and verified with the experimental results obtained from embedded thermocouples.

For these angles temperature history from the model is recorded and verified with the experimental results obtained from embedded thermocouples. Differences in the temperatures for the different angles cause different heat treatment of the deposited material and the material in the vicinity of the deposited material, i.e. the heat affected zone. Although the different grain sizes are observed for different angles, the phase constituents at different angles are similar. A new experimental set-up is designed and a 3D part is built.Acknowledgment The authors wish to acknowledge the financial support provided by the THECB, Grants No. 003613-0022001999 and 003613-0016-2001, NSF, Grants No. DMI – 9732848 and DMI – 9809198, and the U.S. Department of Education, Grant No. P200A-80806-98.References: (1) Spencer, J.D., Dickens, P.M., and Wykes, C.M., (1998). Rapid prototyping of metal parts by three-dimensionalwelding, ImechE Proc. Part B, Journal of Engineering Manufacture, Vol. 212, pp.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffect of Voronoi Lattice Geometry on the Fatigue Performance of Ti-6Al-4V S.M. Uí Mhurchadha1, S. Marques1, L. Givet1, R. Raghavendra1, 2 1 SEAM Research Centre, School of Engineering, Waterford Institute of Technology, Ireland 2 I-Form Advanced Manufacturing Research Centre, IrelandAbstractThis paper investigates the effect of strut thickness and number of pores on the fatigue performance ofTi-6Al-4V voronoi lattice structures designed with the same part volume. The aim of this study is to establish the variation in high cycle fatigue parameters for constant volume lattice structures designed with various lattice parameters. Voronoi geometries were designed with varying strut thicknesses and number of pores to maintain a constant specimen volume.

Voronoi geometries were designed with varying strut thicknesses and number of pores to maintain a constant specimen volume. The geometries were tested under compressive fatigue conditions at a reversal ratio, R, of 0.1. It was found that the strut thickness has a significant influence on the fatigue life of the lattice. An increase in the strut thickness by 100 µm can result in a reduction in fatigue life by up to a factor of 10. The results from this research can influence the design of lattice structures for osteointegration in load- bearing biomedical implant applications. Keywords: High-cycle fatigue, Metamaterials, Laser Powder Bed Fusion, Fatigue lifeIntroductionAdditive manufacturing (AM) allows for the manufacture of complex lattice structures by using a layer- by-layer approach to manufacturing parts. Lattice structures are also known as metamaterials, which are structures that are designed so that their properties are primarily geometrically-driven.

Lattice structures are also known as metamaterials, which are structures that are designed so that their properties are primarily geometrically-driven. For biomedical applications, these lattice structures can be used in orthopaedic implants to promote osseointegration, as well as mimicking the mechanical properties of human bone. Typically, for orthopaedic implants, the pore size range of these lattice structures should be in the range of 150 to 600 µm [1]. The reported porosity of trabecular bone ranges from 50 to 90 % [2, 3]. Biocompatible materials, such as the Titanium alloy, Ti-6Al-4V, are used for the manufacture of orthopaedic implants.Researchers have used various methods of generating bio-mimicking lattice structures, such as repeated cellular units [4, 5], TPMS (triply periodic minimal surface) [6, 7], stochastic lattices based on a Poisson disk algorithm [4, 8, 9] and Voronoi tessellation [10, 11]. This research focuses on the Voronoi tessellation method.

This research focuses on the Voronoi tessellation method. A Voronoi lattice is a 3D isotropic porous interconnected model that can be designed to mimic the physical properties of bone, such as trabecular thickness, trabecular separation, trabecular number, bone volume to total volume ratio and bone surface to bone volume ratio, for example [11].Load-bearing implants in the body undergo cyclic compressive loading, which can result in fatigue damage in extreme cases. ASTM [12] defines fatigue as “the progressive, localised permanent structural change occurring in a material subjected to conditions which produce fluctuating stresses and strains at some point or points which may culminate as cracks”. Under fatigue loading, failure of materials can be observed at loads less than the monotonic yield strength of the material. Fatigue can be categorised into either high-cycle fatigue (HCF) or low-cycle fatigue (LCF), depending on the number of cycles to failure, Nf.

Fatigue can be categorised into either high-cycle fatigue (HCF) or low-cycle fatigue (LCF), depending on the number of cycles to failure, Nf. The total life, Nf, is represented as the sum of the number of cycles to crack initiation, Ni, and the number of cycles for the crack to propagate to a detectable or critical length, NP, where Nf = Ni + NP. Wöhler [13, 14] was the first to carry out a systematic study on fatigue. A relationship between stress and fatigue life was observed; the stress range has a greater influence on fatigue life than the peak stress [15]. From this work, S-N (stress vs number of cycles to failure) curves were proposed to estimate fatigue lives; this S-N curve method is still used today. Since then, many researchers have developed fatigue failure criterion based on stress [16, 17, 18] or strain [19, 20] conditions.Recently, researchers have investigated the fatigue behaviour of metallic metal-biomaterials.

The primary loading conditions on orthopaedic implants are compression and bending. Therefore, many researchers, e.g. [21, 22, 23, 24, 25, 26], have investigated the compressive fatigue loading of bone-mimicking bio-metamaterials, establishing S-N curves for the metamaterials.This paper investigates the effect of the geometry lattice parameters on the fatigue behaviour of Ti-6Al- 4V Voronoi lattices manufactured via Laser Powder Bed Fusion (L-PBF). To this end, lattice structures are13designed to have the same volume, i.e. the same lattice density, but different lattice parameters, such as strut thickness and number of pores. Two lattice geometries are designed with strut thickness differing by 100 µm. These lattices are then tested in compressive fatigue under high cycle fatigue loading conditions.

Two lattice geometries are designed with strut thickness differing by 100 µm. These lattices are then tested in compressive fatigue under high cycle fatigue loading conditions. S-N curves are derived from the experimental results and high-cycle fatigue parameters calculated.MethodologyDesign and manufacture of lattice specimenVoronoi lattice structures were designed using nTopology software [27]. The test samples were designed based on ISO 13314 [28] standard for the compression testing of porous and cellular metals. Cylindrical specimen of diameter 10 mm and height of 15 mm [28] were used for the two lattice geometries, presented in Figure 1, were designed and used in this study. The density of the lattices was design to be approximately 20 %. Often the density of a porous structure is predefined, for example to match the density of a particular bone type.

The density of the lattices was design to be approximately 20 %. Often the density of a porous structure is predefined, for example to match the density of a particular bone type. Therefore, for this study a constant density was used for the design to investigate the effect of the geometry variables that result in the same density. The design parameters for the two geometries are presented in Figure 1.Figure 1. Two Voronoi lattice geometries used in this study, Geometry 1 and Geometry 2, with the same part density, but different strut thicknesses and number of pores.Ten of each lattice structure presented in Figure 1 were manufactured via L-PBF on an EOS M100 machine. The material used was Grade 23 Ti-6Al-4V ELI (Extra Low Interstitials) supplied by Carpenter Additive [29]. The laser spot size of the EOS M100 machine was approximately 40 µm. The build chamber was maintained at 40 ⁰C and under an inert Argon atmosphere throughout the manufacturing process.

The laser spot size of the EOS M100 machine was approximately 40 µm. The build chamber was maintained at 40 ⁰C and under an inert Argon atmosphere throughout the manufacturing process. The specimen were removed from the build plate using Wire EDM. Powder was removed for the lattice structures using a combination of dry ice blasting and ultrasonic cleaning.Characterisation of latticesThe volume of the manufactured lattices was measured using Helium gas pycnometry on a Micrometics Accupyc 1340 gas pycnometer. This is a volume measurement methodology based on Boyle’s Law. The specimen is placed in a chamber and Helium (He) is pumped into the chamber to a known a set pressure. This volume of He is then released into a second chamber of known volume and the pressure of the gas in this chamber is measured. From this, the volume of the lattice structure is calculated. The pycnometry tests were repeated five times per sample and then averaged.

From this, the volume of the lattice structure is calculated. The pycnometry tests were repeated five times per sample and then averaged. He gas was used as it has low density and can penetrate small surface pores to get an accurate measurement of the volume. The mass of each specimen was also measured.14Fatigue testingCompressive fatigue tests were conducted on the specimen using an Instron Electropulse 3 kN testing machine at a frequency of 15 Hz [22]. A reversal load ratio, R, of 0.1 was used for the loading conditions, ensuring that the specimen were under compressive loading throughout all tests. R is given by:𝑅 =𝐹𝑚𝑖𝑛 𝐹𝑚𝑎𝑥where Fmin and Fmax are the minimum and maximum forces for each fatigue cycle. The maximum and minimum applied forces for each loading condition is presented in Table. The fatigue tests were stopped after 1 million cycles if the specimen did not fail.Table 1.

The maximum and minimum applied forces for each loading condition is presented in Table. The fatigue tests were stopped after 1 million cycles if the specimen did not fail.Table 1. Loading conditions for compressive fatigue testing of lattices.Loading Condition LC1 LC2 LC3 LC4 LC5Fmax (N) -2852 -2333 -1815 -1296 -778Fmin (N) -285.2 -233.3 -181.5 -129.6 -78.8Fatigue material modelFrom the fatigue test results, S-N curves for each lattice geometry were correlated. The stress was calculated based on the force per unit area, which was calculated based on the average cross-sectional area per unit height of the lattice. A comparison between the designed geometry and actual geometry was investigated.Basquin [17] developed the stress-life approach, by observing a log-log relationship between stressamplitude and the fatigue life for high-cycle fatigue (HCF).

The Basquin equation is given by: ∆𝜎 2𝑏′(2𝑁𝑓)= 𝜎𝑓′ is the fatigue strength coefficient and b is the fatigue strength exponent. Fatiguewhere Δσ is the stress range, 𝜎𝑓 parameters for the two lattice geometries are compared.Results and discussionCharacterisation of manufactured latticesThe measured volumes for the two lattice geometries are presented in Figure 2. There is a large difference between the designed and manufactured volumes. This is associated with the partially sintered particles on the struts of the lattice, as well as the surface roughness of these struts. This has a larger effect of the lattice volume for Geometry 2, due to the larger surface area associated with the thinner struts combined with the larger number of pores. The strut thicknesses of the two geometries are shown in the SEM images presented in Figure 3. The average strut thickness of the as-build Geometry 1 is 350 µm, which is the same as the designed strut thickness.

The average strut thickness of the as-build Geometry 1 is 350 µm, which is the same as the designed strut thickness. However, the average strut thickness for Geometry 2 is 290 µm, which is 40 µm larger than the designed thickness. This can be attributed to the accuracy of the laser in the L-PBF for hatch strategies for smaller parts.15(1)(2)Figure 2. Manufactured volumes and densities of the two voronoi geometries compared to the design criteria.(a)(b)Figure 3. SEM images of (a) Geometry 1 with an average strut thickness of 350 µm, and (b) Geometry 2 with average strut thickness of 290 µm.Fatigue life of latticesFigure presents the force-displacement loading loops for both geometries, for the maximum loading conditions, i.e. Fmax of 2.852 kN, at for initial cycle, mid-cycle and final cycle of the compressive fatigue test.

Fmax of 2.852 kN, at for initial cycle, mid-cycle and final cycle of the compressive fatigue test. These loops confirm that no plasticity occurred in the samples during testing and therefore, all fatigue tests conducted were under high cycle fatigue loading conditions. However, it is possible that some localised plasticity occurred at highly localised stress concentrations within the lattice structure. Therefore, Figure demonstrates that the dynamic mechanical behaviour of the metamaterial was elastic and therefore, tested in high cycle fatigue.16(a)(b)Figure 4. Force displacement loops for the initial cycle, a mid-cycle and the final cycle for LC1 compressive fatigue loading conditions on (a) Geometry 1, and (b) Geometry 2.Summaries of the minimum and maximum stresses for each loading condition on the two geometries are presented in Figure . The stresses presented here are based on the designed cross-sectional area of the specimen.

The stresses presented here are based on the designed cross-sectional area of the specimen. This is to highlight the requirement to determine the actual cross-sectional area of the lattices after manufacturing. For Geometry 1 (Figure (a)), the test reached 1 million cycles for LC3, as presented in Table. Whereas, Geometry 2 (Figure (b)) reached 1 million cycles for LC5 as presented in Table. For both geometries, a life of over 106 cycles was determined to be infinite cycles. A considerable difference can be seen between the two geometries.(a)(b)Figure 5. Minimum and maximum stress evolution, as calculated using the designed cross-sectional area, for each load case, as presented in Table, for (a) Geometry 1, and (b) Geometry 2.Figure presents the evolutions of the minimum and maximum actual stresses, based on the measured L- PBF lattices, for each loading condition on the two geometries.

As expected from the measured volume results presented in Figure 2, the actual stresses are lower compared to those presented in Figure 5. Geometry 2 has a significantly lower load-bearing capacity compared to Geometry 1. For Geometry 2, LC2 had a maximum compressive stress range (Δσ/2 = (|σmax - σmin|)/2) of 33 MPa, similar to the compressive range stress of 39 MPa experienced by Geometry 1 for LC3. However, Geometry 1 test was considered to run infinitely i.e. greater than 106 cycles, whereas, Geometry 2 failed after 4.8 × 104 cycles. This is a considerable difference in the dynamic load-bearing capability of the two structures under similar loading conditions. Despite Geometry 2 being manufactured to have a larger cross-sectional area, it has a lower fatigue load carrying capabilities compared to Geometry 1.

Despite Geometry 2 being manufactured to have a larger cross-sectional area, it has a lower fatigue load carrying capabilities compared to Geometry 1. This can be attributed to the increased surface area of Geometry 2, which provides more crack initiation points, combined with the thinner struts, meaning that once a crack initiations, it requires fewer cycles to propagate to failure of the structure.17(a)(b)Figure 6. Minimum and maximum stress evolution, as calculated using the actual cross-sectional area, for each load case, as presented in Table, for (a) Geometry 1, and (b) Geometry 2.Figure 7 presents the S-N curves for Geometry 1 and Geometry 2, as calculated for the design stress (Figure 7 (a)) and the actual stress (Figure (b)) for each sample. For both the design and the actual stress graphs, a significant difference between the two geometries can be observed. As seen in Figures 5 and 5, Geometry 1 have a greater dynamic load-bearing capacity compared to Geometry 2.

As seen in Figures 5 and 5, Geometry 1 have a greater dynamic load-bearing capacity compared to Geometry 2. Critically, there is a significant difference between the S-N curves, as calculated using the designed cross-sectional area and as calculated using the actual cross-sectional area. For example, for Geometry 2 and at a stress range of 40 MPa, the fatigue life using the actual cross-sectional area S-N curve is in the range of 1.5 × 104; whereas, if the designed cross-sectional area S-N curve is used, the fatigue life is in the range of 5.6 × 105. This represents a factor greater than 10 between the two S-N curves, highlighting the requirements to measure the actual manufactured lattice structure rather relying solely on the design data.Overall, Geometry 2 has a considerably lower fatigue life, compared to Geometry 1. This can be attributed to a few different factors.

This can be attributed to a few different factors. Firstly, the thinner struts in Geometry 2 results in fewer microstructural grains in the cross-section of the strut, compared to Geometry 1. Grain boundaries have been shown to hinder crack propagation of fatigue cracks, resulting in increased cycles for crack propagation [30]. Therefore, the greater the number of grain boundaries in a cross-section of a lattice strut, the greater the number of cycles for crack propagation. Furthermore, the larger number of pores in the lattice of Geometry 2 means that there is a larger probability of a stress concentration occurring. These stress concentration also have a larger effect on the structure of Geometry 2 due to the thinner struts. Finally, as presented in Figures 2 and 3, Geometry 2 has a larger surface area compared to Geometry 1. This provides more area for crack initiation points, which are further exacerbated by the inherent surface roughness from L-PBF.(a)(b)Figure 7.

This provides more area for crack initiation points, which are further exacerbated by the inherent surface roughness from L-PBF.(a)(b)Figure 7. S-N curves for Geometry 1 and Geometry 2 for (a) the design stress for each sample, and (b) the actual stress for each samples.18Fatigue coefficients′ and b) can be estimated from monotonic test data using the universal slopes method [17, 31]. This method gives estimated fatigue coefficients based on monotonic tensile test results for ′ = 1.9σUTS, and b = −0.12. This method has been used to provide a baseline ultimate tensile strength, σUTS, e.g. 𝜎𝑓 for the high-cycle fatigue parameters for L-PBF Ti-6Al-4V, presented in Table 2 [30].Basquin parameters (𝜎𝑓Basquin high-cycle fatigue parameters for Geometries 1 and 2 were obtained from the HCF test data presented above. Table 2 also presents the Basquin high-cycle fatigue parameters for Geometries 1 and 2, based on the stress calculated using the designed cross-sectional area.

Table 2 also presents the Basquin high-cycle fatigue parameters for Geometries 1 and 2, based on the stress calculated using the designed cross-sectional area. The fatigue strength exponent, b, is about twice ′, for both that of the baseline for L-PBF Ti-6Al-4V, shown in Table 2. The fatigue strength coefficients, 𝜎𝑓 geometries are lower than that of the baseline for L-PBF Ti-6Al-4V. The Basquin high-cycle fatigue parameters for Geometries 1 and 2, based on the stress calculated using the actual cross-sectional area of the parts are also presented. The fatigue strength exponent, b, for each geometry is the same as those presented in Table 2. However, ′, for both geometries are lower than those calculated for the designed cross- the fatigue strength coefficients, 𝜎𝑓 sectional area of the specimen.Table 2.

However, ′, for both geometries are lower than those calculated for the designed cross- the fatigue strength coefficients, 𝜎𝑓 sectional area of the specimen.Table 2. Baseline high-cycle fatigue parameters for L-PBF Ti-6Al-4V, calculated using the Universal Slopes Method [31] and for Geometries 1 and 2, calculated using the designed and manufactured cross-sectional areas of the structures.Universal slopes method [30]Based on designed geometryBased on manufactured geometryMaterial Ti-6Al-4V Geometry 1 Geometry 2 Geometry 1 Geometry 2′ (MPa) 1710 1360 908 975 470𝝈𝒇b -0.12 -0.21 -0.22 -0.21 -0.22Figure presents HCF experimental test results with Basquin fit for Geometries 1 and 2, from Table 2, compared to the baseline Basquin equation for L-PBF Ti-6Al-4V, for the stress range calculated using the designed cross- sectional areas (Figure (a)) and the actual cross-sectional areas for the specimen (Figure (b)).

There is a good fit between the experimental data and the Basquin equations.(a)(b)Figure 8. HCF experimental test results with Basquin fit for Geometries 1 and 2, compared to the baseline Basquin equation for L-PBF Ti-6Al-4V, for the stress range calculated using (a) the designed cross-sectional areas, and (b) the actual cross-sectional areas for the specimen.DiscussionThe use of the Basquin equation alone assumes that the fatigue loading on the structures is uniaxial, where the largest principal stress is much greater than the other principal stresses. Due to the design of the structures, a19multiaxial fatigue approach may be more accurate for the prediction of fatigue in complex structures. In lattice structures, multiaxial fatigue can be further complicated by the potential for non-proportional loading.

In lattice structures, multiaxial fatigue can be further complicated by the potential for non-proportional loading. Under proportional loading conditions the principal stresses remain in constant proportion to each other, and hence the angle of the principal axes remains constant throughout the fatigue cycle. Non-proportional or out-of-phase loading results in extra cyclic hardening due to the rotation of principal axes. The critical plane approach was developed from the work of Brown and Miller [32]. It is based on the physical observation that fatigue cracks initiate and grow on certain planes within a material, where the crack growth and orientation depend on the normal stresses and/or shear stresses and strains on these planes [33]. The planes on which most fatigue damage occurs are called critical planes.

The planes on which most fatigue damage occurs are called critical planes. This approach overcame the limitations of the Brown and Miller [32] approach with respect to the increased damage associated with out-of-plane loading.ConclusionsThis paper presented an investigation into the effect of Voronoi lattice geometry on the fatigue performance of Ti-6Al-4V. Two different lattice structures with the same overall lattice density were design in nTopology. The strut thicknesses of the lattices differed by 100 µm. Therefore, the number of pores also different, to maintain a constant volume between the designs. These designs were manufactured via L-PBF and the differences between the designed and manufactured part correlated. The two lattice geometries were dynamically tested under compressive fatigue loading and fatigue parameters defined.

The two lattice geometries were dynamically tested under compressive fatigue loading and fatigue parameters defined. The primary findings of the paper can be concluded as follows:A lattice geometry design with a larger surface area results in a larger discrepancy between the designed volume and the actual L-PBF manufactured volume. This is due to the adherence of particles to the surface of the lattice struts. The increased surface area resulted in more crack initiation points due to the inherent surface roughness of L-PBF parts.The high-cycle fatigue load-bearing capacity of the lattices structure is highly dependent on the lattice geometry, in particular, the strut thickness. The lattice structure with thinner struts exhibited considerably lower fatigue strength compared to the thicker lattice with larger strut thickness.A reduction in lattice strut thickness can result in a reduction in the fatigue life of the lattice by a factor of10.

It is essential to measure the difference between the designed and actual cross-sectional area of the lattice structures. Failure to do so could result in the over prediction of fatigue life of a component.Finally, this paper has outlined the high cycle fatigue performance considering the lattices structure as a metamaterial. However, further in depth studies of the multi-axial fatigue of a lattice structure will provide an understanding into the failure mechanisms. This will require the implementation of critical plane multi-axial fatigue analysis of lattices structures under dynamic loading within a finite element simulation. By understanding the failure mechanisms of the lattice structures, the design of lattices can be tuned to prolong the fatigue life of the structures.AcknowledgementsThe authors wish to acknowledge 3DWIT for the Wire EDM of the samples from the build plate.References[1] C.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEffect of Wall Thickness and Build Quality on the Compressive Properties of 304L Thin-walled Structures Fabricated by SLM Myranda Spratt1, Sudharshan Anandan2, Rafid M. Hussein2, Joseph W. Newkirk1*, K. Chandrashekhara2, Heath Misak3, and Michael Walker31Materials Science and Engineering Department2Department of Mechanical and Aerospace EngineeringMissouri University of Science and Technology, Rolla MO 654093Spirit AeroSystems, Inc, Wichita, KS 67210Author for Correspondence (jnewkirk@mst.edu)Keywords: Steel, SLM, LatticeAbstractThe specific strength of lightweight lattice structures built with SLM is of interest to the aerospace industry. Honeycombs were manufactured with increasing wall thicknesses (which increases density) and tested under compression.

Honeycombs were manufactured with increasing wall thicknesses (which increases density) and tested under compression. The optimal strength to density ratio was determined from the resulting data. The build quality was also evaluated to determine how/if the results were influenced by the specimen quality. Differences between the nominal and as-built geometry were identified, but considered to be minimal. Microstructural evaluation of the specimens revealed a possible dependence on the ‘border scan’ properties, as the thickness of the specimens was such that the board scan made up most of the part. This work was used to validate the results of a finite element analysis of this geometry.IntroductionLow density, high strength materials are used in a number of applications such as theaerospace [1] and automotive industries [2] and as lightweight protective equipment [3].

Carbon fiber reinforced polymers have been used recently in applications where weight saving is needed, such as in newer airplanes, but are limited by temperature. [4] Figure I is a material design chart [5] that shows the density-strength relationship of many composite and monolithic materials. As this chart shows, the specific strengths of most materials follow a linear trend on a log-log graph. Efforts to improve specific strength are on-going and often involve optimizing complex structures.708Figure I: Ashby material design chart for strength and density, where materials are plotted to show general trends. [5]Lightweight structure design can be made simpler using finite element analysis (FEA) to create optimal structures, which can then be additively manufactured. Structures with complex geometries can be computationally intensive to optimize, so a unit cell approach can be used.

Structures with complex geometries can be computationally intensive to optimize, so a unit cell approach can be used. A unit cell approach is typified by designing a unit cell which can be repeated in 2 or 3 dimensions to create the entire structure. A representative volume element (RVE) can then be selected to undergo FEA. Figure II shows three samples of unit cells – a 2D and 3D strut based design and a minimal surface design. The hexagonal honeycomb (the 2D strut based design) was selected for this experiment as it is the simplest of the three structures. Using mechanical property data for the material, a researcher can build a computationally-friendly RVE from the unit cell to represent the entire structure. This can then be optimized by FEA for the necessary application.

This can then be optimized by FEA for the necessary application. In this process, all that is needed are the mechanical properties data for the SLM manufactured material and experimental validation of the results after FEA.Digital models of components can be produced as a result of computational design. With only a few limitations, the digital model can then be fabricated with additive manufacturing. Additive manufacturing (AM) is a layer-by-layer process in which a digital component can be ‘sliced’ into layers and printed. There are several challenges to overcome when fabricating components with AM. One challenge is that the digital component, upon which any modeling is based, will not be perfectly reproduced with AM. Deviations from the model can include surface roughness, warpage due to residual stress, printing defects, and more. [6] AM-built components can also be anisotropic due to the layered material.

Deviations from the model can include surface roughness, warpage due to residual stress, printing defects, and more. [6] AM-built components can also be anisotropic due to the layered material. [7] This runs contrary to most material709B Figure II: Unit cell designs where A) is a 2D honeycomb, B) is a 3D body centered cubic strut design, and C) is a 3D gyroid minimal surface design.CAmodels which assume the properties are isotropic. Other problems arise depending on the exact type of processing. Selective laser melting (SLM), for example, cannot print enclosed features without trapping powder within the component, which can be a major design limitation.In this paper, the difference between the digital model of a honeycomb structure and the as-built component using simple physical measurements was determined. This difference can be used to inform the FEA model, the subject of a future paper.

This difference can be used to inform the FEA model, the subject of a future paper. The compressive strength of honeycombs of increasing wall thickness was determined and compared to the strength to density of other materials. Finally, microstructure evaluation of the honeycombs after compression was used to aid in the failure analysis.Materials and MethodsThe honeycomb specimens were printed in a selective laser melting machine – a Renishaw AM250. 304L powder from LPW was used to build the specimens, in an argon atmosphere. The machine parameters were optimized for 304L and included a border scan optimized to decrease surface roughness and an internal scan optimized to increase density. The specimens were printed according to the specifications in Table I. They were removed from the build plate via wire electrical discharge machining (EDM), and cleaned with an ultrasonic bath to remove any remaining loose powder. Physical dimensions were measured with calipers.

Physical dimensions were measured with calipers. The density was determined using Archimedes’ method, with 95% isopropyl alcohol as the immersion liquid.The specimens were compression tested using ASTM C365 at a strain rate of 3 mm/min. Ineach sample set, one specimen was randomly selected to not be tested. These specimens were sectioned, polished and etched to optically image the microstructure. After compression testing, other specimens were also selected for imaging. The specimens were electrolytically etched in a 60:40 nitric acid: water solution.710Table I: Sample dimensions as specified by the digital model.Sample Designations: Wall Thickness (mm) Cell Size (mm) Height (Z) Width (W) Length (L)H1 0.2 3.97 25.4 25.15 26.67H2 0.3 3.97 25.4 25.73 27.28H3 0.4 3.97 25.4 24.38 26.67H4 0.5 3.97 25.4 20.83 22.86Results and DiscussionPrint accuracy was measured by comparing the CAD model of the honeycomb to the printed sample.

Figure III shows a side-by-side view of the two. The dimensions measured were height, width, and length of the overall specimen, the cell size, and the wall thickness. Of all the measurements, only the wall thickness showed a statistically significant difference between the CAD model and the as-built samples. The details can be seen in Table II. The wall thickness was consistently larger than the CAD model, between 90 and 140 microns.

The details can be seen in Table II. The wall thickness was consistently larger than the CAD model, between 90 and 140 microns. This indicates that in the range of wall thicknesses tested, the accuracy of the Renishaw was generally good.WidthCell SizeLengthAHeightBWall ThicknessFigure III: Side-by-side view of the CAD model (A) and as-built (B) honeycombs.Table II: Measured wall thickness compared to the digital model, with the standard deviation on each average value.Sample Designations: Nominal Wall Thickness (mm) Measured Wall Thickness (mm) 0.3±0.01H1 0.2H2 0.3 0.44±0.02H3 0.4 0.52±0.02H4 0.5 0.59±0.02711The difference between the nominal and measured wall thicknesses appears to be causedby a similar source across the samples. The deviation was attributed to the presence of powder particles sintered to the surface of the specimens.

The deviation was attributed to the presence of powder particles sintered to the surface of the specimens. The size of the 304L powder was between 45 and 65 microns, which (when doubled to account for each wall of the specimens) corresponds to the increase in wall thickness from the CAD model to the as-built samples. These particles occur during the melting stage of the process. Particles can be heated due to proximity to the melt pool and can then weld to it. In a part where the surface area is very high (such as in these honeycomb samples), there is concern that these particles could detrimentally affect the mechanical properties of the component.Investigation of the microstructure before and after compression testing reveals several key features about these structures. First, the surface particles sintered to the honeycomb walls are easily visible and correspond to the 45 to 65 micron range suggested above. Some of these particles are circled in Figure IV.A.

First, the surface particles sintered to the honeycomb walls are easily visible and correspond to the 45 to 65 micron range suggested above. Some of these particles are circled in Figure IV.A . Second, the melt pools and track boundaries are outlined in white after etching, with a dark feature within the bands. A close inspection of the dark features (Figure IV.B) finds that it is a cellular structure composed of dark and white features. Third, Figure V shows a honeycomb strut after compression testing. From this, and macro observation of the honeycomb, it was found that the honeycomb failed by local buckling near the center of the structure. Finally, there is very little porosity visible in these images, which indicates a high material density.

Finally, there is very little porosity visible in these images, which indicates a high material density. This was confirmed by Archimedes density testing, which found that, on average, the density of the material was above 98% of theoretical.ABFigure IV: Honeycomb microstructures where A) shows a triple point and B) shows the cellular features. The blue circle shows particles welded to the walls, the white arrows show the build direction.712Figure V: A honeycomb strut that buckled after compression testing.Table III shows the results of compression testing for each sample set. The density of thehoneycombs increased from ~1g/cc to ~2g/cc as shown in Table III. The ultimate and yield strengths of the honeycombs increased as the density of the honeycombs increased, as expected. However, the specific yield stresses did not appear to increase as density increased, and though the specific ultimate strength did.

However, the specific yield stresses did not appear to increase as density increased, and though the specific ultimate strength did. The yield stresses of the honeycombs were plotted against the densities of the honeycombs on the material design chart in Figure VI [5]. The honeycombs yielded in the area marked as the polymers and elastomers region of the chart. They performed very similarly to most common plastics (PMMA, PEEK, PA, etc).

The honeycombs yielded in the area marked as the polymers and elastomers region of the chart. They performed very similarly to most common plastics (PMMA, PEEK, PA, etc). The honeycombs could also be considered as an extension of the foam region as they follow a similar trend.Table III: The results of compression testing the honeycomb samples, with the standard deviation on each average value.Sample Designations:H1H2H3H4Density (g/cm3=Mg/m3)1.05±01.46±0.131.67±0.021.97±0.01Yield Stress (MPa)45.4±2.254.8±2.675±3.883.3±0.2Ultimate Stress (MPa)91.9±9.1132.8±7.8194±2.4267.1±6.2Specific Yield (MPa/g/cm3)43.2±2.137.6±1.844.9±2.342.2±0.1Specific Ultimate (MPa/g/cm3)87.3±8.581.8±5.5116.5±2.1134.4±3.2713Figure VI: Strength-density chart including the yield stress data from this study (purple boxes).

[5]ConclusionsIn conclusion, this study found that the honeycomb parts built with the Renishaw AM250 closely matched with the digital model. The only exception to this were particles that welded to the surface of the honeycombs during printing, which inflated the wall thickness of the honeycombs. Microstructural evaluation of the honeycombs found that localized buckling of the cell walls appears to have caused the honeycomb failure. The yield stress of the compressed samples, when graphed against other materials of the same density, was comparable to common structural plastics.AcknowledgementsSupport from Center for Aerospace Manufacturing Technologies (CAMT) at Missouri University of Science and Technology is gratefully acknowledged. The authors would also like to acknowledge Honeywell Federal Manufacturing and Technology for use of the Renishaw machine.714References[1] M. Standridge, Aerospace materials — past, present, and future, Aerosp. Manuf. Des. (2014).

FECT OF FREE-EDGES ON MELT POOL GEOMETRY AND SOLIDIFICATION MICROSTRUCTURE IN BEAM-BASED FABRICATION OF THIN-WALL STRUCTURESJoy E. Davis, Nathan W. Klingbeil and Srikanth Bontha Department of Mechanical and Materials Engineering Wright State University Dayton, OH 45435AbstractThe success of both laser and electron beam-based fabrication processes for additive manufac- turing and repair applications requires the ability to control melt pool geometry while still main- taining a consistent and desirable microstructure. To this end, previous work by the authors has employed point-heat source solutions to investigate the effects of process variables (beam power and velocity) on melt pool geometry and solidiﬁcation microstructure (grain size and morphol- ogy) in beam-based fabrication of thin-wall structures. However, these results were limited to steady-state conditions away from free-edges.

However, these results were limited to steady-state conditions away from free-edges. The current work extends the approach to investi- gate transient behavior in the vicinity of a free-edge.IntroductionBoth laser and electron beam-based manufacturing are under consideration for aerospace appli- cations, for which the ability to obtain a consistent and desirable microstructure is of critical con- cern [1–5]. Speciﬁc applications for these processes include fabrication of complex parts, addition of features to existing components and repair of damage or defects. While beam-based additive manufacturing offers the promise of increased efﬁciency and ﬂexibility compared to conventional manufacturing, widespread commercialization of these processes requires the ability to predict and control melt pool size and solidiﬁcation microstructure.

In particular, the mechanical properties of the deposited material are strongly dependent on the solidiﬁcation microstructure (grain size and morphology), which is controlled by the thermal conditions at the onset of solidiﬁcation.The analysis presented herein is based on the 2-D Rosenthal solution for a moving point heat source [6]. The Rosenthal solution was ﬁrst applied to laser based deposition processes by Dykhuizen and Dobranich [7, 8]. The subsequent identiﬁcation of dimensionless process variables by Vasinonta et al. [9, 10] led to the development of thermal process maps for predicting trends in solidiﬁcation microstructure by Bontha et al. [5,11–14]. The Rosenthal solution is applicable for a semi-inﬁnite wall, however it does not address the response in the vicinity of a free-edge. Through direct thermal imaging of the melt pool, Rangaswamy et al. [15] observed that an increase in melt pool size occurs upon approaching a free-edge, without any change in process variables.

Through direct thermal imaging of the melt pool, Rangaswamy et al. [15] observed that an increase in melt pool size occurs upon approaching a free-edge, without any change in process variables. Ag- garangsi et al. [16] developed thermo-mechanical ﬁnite element models to understand and control the increase in melt pool size in the vicinity of a free edge. Nonlinear thermo-mechanical ﬁnite element modeling of laser deposition processes has also been used to predict steady-state melt pool behavior and residual stress [9, 10, 17–19], as well as to establish the utility of the Rosenthal solution for predicting trends in solidiﬁcation microstructure [5, 11–14].In this work, the effect of free-edges on melt pool geometry and solidiﬁcation microstructure is investigated for 2-D thin-wall geometries. Through superposition of the Rosenthal solution,a closed-form solution for the temperature distribution in the vicinity of a free-edge is obtained.

Through superposition of the Rosenthal solution,a closed-form solution for the temperature distribution in the vicinity of a free-edge is obtained. Melt pool contours are plotted as a function of distance from the free-edge for both small-scale (e.g., LENSTM) and large-scale (higher power) processes. In addition, cooling rates and thermal gradients at the onset of solidiﬁcation are numerically extracted throughout the depth of the melt pool, and are subsequently plotted on solidiﬁcation maps for predicting trends in grain size and morphology for Ti-6Al-4V. Finally, ﬁnite element analysis is performed to verify the validity of the closed-form solution, as well as to investigate solidiﬁcation behavior when the beam reaches the edge and is turned off.2-D Rosenthal SolutionThis study considers the 2-D thin-wall geometry of Fig. 1, in which the process variables of interest are the absorbed beam power αQ and velocity V.

1, in which the process variables of interest are the absorbed beam power αQ and velocity V . It is assumed that the height h and length L are sufﬁciently large such that the steady-state Rosenthal solution for a moving point heat source applies.αQVxozohLbFigure 1: 2-D GeometryFor the geometry of Fig. 1, the temperature of the wall relative to the position of the point heatsource has been given in dimensionless form by Vasinonta et al. [9, 10] as(cid:2)(cid:3)(cid:4)T = e−x0K0x2 0+ z2 0,where K0 is the modiﬁed Bessel function of the second kind, order zero. The dimensionless vari- ables in eq. (1) are deﬁned in terms of the absorbed beam power αQ and velocity V asT = T − T0αQ πkb,x0 = x02k ρcVandz0 = z02k ρcV.In the above normalizations, T is the temperature at a location (x0,z0) relative to the moving point heat source and T0 is the initial temperature of the wall.

The relative coordinates (x0,z0) are related(1)(2)to ﬁxed spatial coordinates (x, z) at any time t as (x0,z0)=(x − V t,z), where V is the velocity of the beam. Also in eq. (4), the thermophysical properties ρ, c and k are the density, speciﬁc heat and thermal conductivity of the material, which are assumed to be temperature-independent. In the subsequent sections, results are presented for both small-scale and large-scale processes, which are characterized in terms of the dimensionless melting temperatureT m = Tm − T0αQ πkb.Closed-Form Solution in Vicinity of Free-EdgeFor the case of temperature-independent properties, a closed-form solution for the temperature distribution in the vicinity of the free-edge can be modeled by superposition of two point-heat sources approaching one another, with the line of symmetry representing the free-edge. This is because the symmetry condition is equivalent to thermal insulation, which is essentially the case at the free-edge.

This is because the symmetry condition is equivalent to thermal insulation, which is essentially the case at the free-edge. This of course assumes that convection and radiation are negligible compared to conduction, which is generally well accepted for beam-based deposition processes.Figure 2: Free-Edge RepresentationAs illustrated in Fig. 2, the dimensionless distance from the free-edge is represented by the parameter a. The normalized distance a/l is used to relate the dimensionless Rosenthal melt pool l to the distance from the free-edge, wherel = l2k ρcVand¯a = a2k ρcV.It is important to note that a/l has the same value as a/l, since the nondimensionalization cancels out. As evident from Fig. 2, the effect of the second heat source can be obtained by replacing x0with the quantity (2a-x0) in the Rosenthal solution of eq. (1).

For each case, the melt pool geometry is plotted a function of distance from the free-edge. In addition, the dimensionless cooling rate and thermal gradient are numerically extracted and evaluated along the trailing edge of the melt pool. Results are subsequently plotted on solidiﬁcation maps for Ti-6Al-4V. In so doing, the thermophysical properties are assumed constant at the melting temperature Tm = 1654◦C, while the absorption coefﬁcient is assumed to be α = 0.35. Finally, the wall thickness is assumed constant at b = 2.26 mm and the velocity at V = 8.47 mm/s, which are in keeping with the values used in the authors’ prior work [5, 11, 12, 20].(cid:4)⎫ ⎬⎭.(6)(7)(8)(9)Effect of Free-Edge for Small-Scale ProcessesDimensionless melt pool geometries are shown in Fig.3 for the case of T m = 2.88. With the parameters speciﬁed above, this corresponds to a beam power of Q = 350 W, which is typical for LENSTM deposition of Ti-6Al-4V.

With the parameters speciﬁed above, this corresponds to a beam power of Q = 350 W, which is typical for LENSTM deposition of Ti-6Al-4V. The steady-state Rosenthal melt pool is seen on the far left. The melt pool is an almost circular cross section and increases in size as the point heat source approaches the free-edge.Figure 3: Small-Scale (LENSTM) Process Melt Pool Geometry, T m = 2.88(cid:1)T (cid:1)tz 0 z m(a) LENSTM Process Cooling Rate(b) LENSTM Process Thermal GradientFigure 4: Small-Scale (LENSTM) (a) Cooling Rate and (b) Thermal GradientRepresentative results for the dimensionless cooling rate are plotted in Fig. 4 (a) as a function of relative depth within the melt pool for selected values of a/l. The relative depth within the melt pool varies in the range 0 ≤ z0/zm≤ 1, where zm is the maximum depth of the melt pool forthe given value of T m.

The relative depth within the melt pool varies in the range 0 ≤ z0/zm≤ 1, where zm is the maximum depth of the melt pool forthe given value of T m. As seen in the ﬁgure, the cooling rate decreases (and ultimately becomes negative) as the point heat source approaches the edge. This is in keeping with the increase in melt pool size illustrated in Fig.3, which is somewhat analogous to an increase in incident energy. The negative cooling rate signiﬁes heating (as opposed to cooling) of the melt pool boundary, as there is nowhere for the heat to go at the free-edge. Representative results for the dimensionless thermal gradient are shown in Fig. 4 (b). The thermal gradient decreases as the beam approaches the free edge, which is again analogous to an increase in incident energy.

4 (b). The thermal gradient decreases as the beam approaches the free edge, which is again analogous to an increase in incident energy. However, the thermal gradient appears to be somewhat less sensitive to the free edge than the cooling rate, particularly near the surface of the deposit.Effect of Free-Edge on Large-Scale ProcessesFigure 5: Large-Scale Melt Pool Geometry, T m = 0.5(a) Large-Scale Process Cooling Rate(b) Large-Scale Process Thermal GradientFigure 6: Large-Scale (a) Cooling Rate and (b) Thermal GradientDimensionless melt pool geometries are shown in Fig.5 for the case of T m = 0.5. With the parameters previously speciﬁed, this corresponds to a beam power of Q = 2013 W, which is more typical of large-scale processes. Compared to the near-circular cross sections of the small-scale melt pools, the large-scale melt pools have an elongated, surf board-type shape. As the melt pool approaches the free-edge the size does not observably change.

As the melt pool approaches the free-edge the size does not observably change. This is because the trailing edge of the melt pool is so far away from the free-edge that the insulated condition has little effect on the overall melt pool geometry. That said, it is expected that the melt pool geometry would be affected if the beam were to reach and subsequently return from the free-edge, which is a subject of ongoing investigation. The dimensionless cooling rate and thermal gradient for the large-scale process can be seen in Fig. 6. As shown in the ﬁgure, both the cooling rate and thermal gradient are insensitive to the free-edge, which is consistent with the negligible change in melt pool geometries shown in Fig.5.Finite Element AnalysisFinite element analysis has been used to include the nonlinear effects of temperature-dependent properties and latent heat, as well as to investigate the solidiﬁcation behavior when the beam reaches the free-edge and is subsequently turned off.

A representative mesh is shown in Fig. 7, and follows procedures outlined in [5, 13]. The movement of the beam begins at the left edge of the wall, and travels left to right across the wall. The model has three different mesh densities in the x-direction. A coarser mesh is used for the ﬁrst sixth of the model, while the element size in the next sixth of the model is one half of the coarser region element size. Finally, the remaining portion of the model has a ﬁner mesh where the element size is one fourth of the coarser region element size.

Finally, the remaining portion of the model has a ﬁner mesh where the element size is one fourth of the coarser region element size. Results from transient conditions near the free-edge are compared with results from steady-state conditions and hence a higher mesh density is also used at the center of the wall.pFigure 7: Representative 2-D Thermal Finite Element Mesh for Transient Analysis(cid:1)aFigure 8: Illustration of the Increase in Melt Pool Size as the Laser Approaches the Free-Edge of the Wall for T m = 2.88Representative melt pool contours are shown for the case of for T m = 2.88 in Fig. 8. Upon approaching the free edge, there is a considerable increase in melt pool size, which is in keeping with the results of Fig. 3.

8. Upon approaching the free edge, there is a considerable increase in melt pool size, which is in keeping with the results of Fig. 3. The effects of steady-state, transient and the stationary melt pool behavior (when the beam is turned off) on solidiﬁcation microstructure are considered in the next section.Effect of Free-Edge on Grain MorphologyAs discussed in [5, 11, 12], results for solidiﬁcation thermal gradient and cooling rate can be interpreted in the context of a solidiﬁcation map to provide predictions of grain size and morphol- ∂t and thermal gradient G = |∇T|, the ogy in Ti-6Al-4V. Given the solidiﬁcation cooling rate ∂T solidiﬁcation rate R is determined asR = 1 G The expected grain morphology can be predicted as either equiaxed, columnar or mixed by plotting points in G vs.

Given the solidiﬁcation cooling rate ∂T solidiﬁcation rate R is determined asR = 1 G The expected grain morphology can be predicted as either equiaxed, columnar or mixed by plotting points in G vs. R space (i.e., on the “solidiﬁcation map”), which has been previously calibrated for Ti-6Al-4V [21].∂T ∂t.(10)Increasing(cid:1)z (cid:1)z0Increasing grain sizeIncreasing grain sizeIncreasing(a) Small-Scale (LENSTM) Process Solidiﬁcation Map, T m = 2.88(b) Large-Scale Process Solidiﬁcation Map, T m = 0.5Increasing depth (cid:1)z (cid:1)z0Increasing grain size(c) FEA Solidiﬁcation Map, T m = 2.88Figure 9: Solidication Maps for (a) Small-Scale (LENSTM) Processes (b) Large-Scale Processes and (c) Finite Element Results(cid:1)z (cid:1)z0Solidiﬁcation maps showing the effect of free-edges on predicted grain morphology for both small-scale (LENSTM ) and large-scale deposition of thin-wall Ti-6Al-4V deposits are shown in Fig. 9.

9. The values of G and R plotted in Fig. 9 (a) and (b) are extracted from the dimensionless cooling rate and thermal gradient plots of Figs. 4 and 6, with thermophysical properties at 1654◦C and process variables as previously noted. The values of G and R plotted in Fig. 9 (c) are extracted directly from the ﬁnite element analysis, with temperature-dependent properties for Ti-6Al-4V.For the small-scale process of Fig.9 (a), the microstructure remains in the fully columnar region as the free-edge is approached. However, the free-edge acts to decrease both the solidiﬁcation rate and thermal gradient, which causes an increase in grain size. So as the free-edge is approached, results suggest a trend toward columnar grains of increasing size.For the large-scale process of Fig. 9 (b), the solidiﬁcation microstructure does not change as the melt pool approaches the free-edge. This is expected based on the results of Fig. 6.

9 (b), the solidiﬁcation microstructure does not change as the melt pool approaches the free-edge. This is expected based on the results of Fig. 6. For all distances from the free-edge there is a grading of the microstructure throughout the depth of the deposit, with a transition from columnar to mixed/equiaxed microstructure at the surface. This is in keeping with results of the authors’ prior work [14].From the ﬁnite element results of Fig. 9 (c), it is clear that the data points from both the steady- state and transient melt pools fall in the fully columnar region, which is in keeping with the results of Fig. 9 (a). However, results also suggest that a transition to equiaxed microstructure can occur when the beam is turned off at the free-edge (stationary melt pool).Summary and ConclusionsThis study has considered the effect of free-edges on solidiﬁcation microstructure (grain size and morphology) in beam-based fabrication of thin-wall structures.

Based on superposition of the Rosenthal solution, a closed-form solution has been presented for the thermal history in the vicinity of a free-edge. Dimensionless cooling rates and thermal gradients have been extracted from the closed-form solution and plotted as a function of normalized distance from the free-edge, for both small-scale (LENSTM) and large-scale processes. In addition, ﬁnite element analysis has been used to include the nonlinear effects of temperature-dependent properties and latent heat, as well as to investigate stationary melt pool behavior when the beam is turned off at the free-edge. Results from both the closed-form and FEA solutions have been plotted on solidiﬁcation maps for predicting grain size and morphology in Ti-6Al-4V. For small-scale processes, results suggest that melt pool geometry may be more sensitive to free-edges than the solidiﬁcation microstructure.

For small-scale processes, results suggest that melt pool geometry may be more sensitive to free-edges than the solidiﬁcation microstructure. This suggests that the control of melt pool size may be possible without affecting microstructure, which is an important result for process developers. Results further suggest that for large-scale processes, both melt pool geometry and microstructure are insensitive to the beam approaching free-edges. A free-edge effect is however anticipated when the beam leaves the free-edge, which is a subject of future work. Finally, ﬁnite element results suggest that a substantial free-edge effect is possible at the end of a ﬁnal pass, when the beam is shut off at the free-edge.

Finally, ﬁnite element results suggest that a substantial free-edge effect is possible at the end of a ﬁnal pass, when the beam is shut off at the free-edge. In particular, results suggest that a transition to equiaxed microstructure may occur even for small-scale processes.AcknowledgmentsThis work has been supported by the National Science Foundation, grant number DMI-0224517, by the joint AFRL/DAGSI Research Program, project number ML-WSU-01-11, and by WrightState University and the Ohio Board of Regents. Any opinions or ﬁndings contained herein are those of the authors, and do not necessarily reﬂect the views of the National Science Foundation, the Ohio Board of Regents or Wright State University.References[1] P. Kobryn, E. Moore, and S. Semiatin, “The Effect of Laser Power and Traverse Speed on Mi- crostructure, Porosity, and Build Height in Laser Deposited Ti-6Al-4V,” Scripta Materialia, vol. 43, pp. 299–305, 2000.[2] F. Arcella and F.

219–226, August 2003.[12] N. Klingbeil, S. Bontha, C. Brown, D. Gaddam, P. Kobryn, H. Fraser, and J. Sears, “Effects of Process Variables and Size Scale on Solidiﬁcation Microstructure in Laser-Based Solid Freeform Fabrication of Ti-6Al-4V,” in Solid Freeform Fabrication Proceedings, (Austin, TX), August 2004.[13] S. Bontha, The Effect of Process Variables on Microstructure in Laser Deposited Materials.PhD thesis, Wright State University, 2006.[14] S. Bontha, N. Klingbeil, P. Kobryn, and H. L. Fraser, “Effects of Process Variables and Size Scale on Solidiﬁcation Microstructure in Beam-Based Fabrication of Bulky 3D Structures,” Materials Science and Engineering A, vol. 513-514, pp. 311–318, 2009.[15] P. Rangaswamy, T. Holden, R. Rogge, and M. Grifﬁth, “Residual Stresses in Components Formed by the Laser Engineered Net Shaping (LENS[registered trademark]) Process,” ournal of Strain Analysis for Engineering Design., vol. 38, pp. 519–527, November 2003.[16] P. Aggarangsi, J.

FECT OF PARTICLE SIZE, BINDER CONTENT AND HEAT TREATMENT ON MECHANICAL PROPERTIES OF 13-93 BIOACTIVE GLASS SCAFFOLDSKrishna C. R. Kolan1, Ming C. Leu1, Gregory E. Hilmas2 and Mariano Velez31Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409 2Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409 3Mo-Sci Corporation, Rolla, MO 65401AbstractParticle size, binder content and the post-processing schedule are important parameters that affect the microstructure, and, hence, the mechanical properties of parts produced using the indirect selective laser sintering process. 13-93 bioactive glass, with mean particle sizes ranging from 10 µm to 44 µm, is mixed with different amounts of stearic acid binder to fabricate green scaffolds.

13-93 bioactive glass, with mean particle sizes ranging from 10 µm to 44 µm, is mixed with different amounts of stearic acid binder to fabricate green scaffolds. Through the design of the post-processing schedule, the time required for post- processing the green scaffolds is reduced from the initial 80 hrs to 12 hrs. The compressive strength varies from 41 MPa for a part with ~60% porosity to 157 MPa for a part with no designed porosity. Several batches of 13-93 scaffolds are soaked in a simulated body fluid for different time intervals ranging from 1 week to 6 weeks. The amount of hydroxyapatite formed and subsequent mechanical properties are provided and discussed.1. IntroductionScaffolds manufactured using biopolymers are bio-inert and do not initiate an active response from the surrounding tissue to bond with the bone by simultaneously supporting tissue regeneration and gene activation. 45S5 glass, discovered by Hench L.L.

45S5 glass, discovered by Hench L.L. in 1969 and popularly known as Bioglass®, was the first glass reported to chemically bond with bone; since then, several compositions have been developed based on this original formulation [1]. The actively researched bioactive materials can be broadly divided into glasses, glass-ceramics and crystalline ceramics. Synthetic hydroxyapatite (HA) is a widely researched ceramic with interesting mechanical properties, but its main limitation lies in its slow resorption and conversion rates. The presence of a heat treatment controlled crystalline phase in glass-ceramic systems increases the mechanical properties of the scaffolds; however, it is reported that the presence of a crystalline phase could reduce the rate of mineralization after implantation [2].

Alternatively, amorphous glasses offer excellent bioactive characteristics and an advanced ability to form an HA layer when exposed to body fluids over anywhere from a few hours to even days, depending on their composition. 13-93 glass, a silicate based bioactive glass, is used in this research work to prepare scaffolds intended for bone repair. A higher amount of SiO2 content in 13-93 glass, compared to 45S5 glass, could slow down the reaction rates upon implantation. However, promising compressive strengths and elastic modulus values for 13-93 makes it a potential candidate for bone scaffolds, particularly for compressive load bearing applications.Two well-known ways of fabricating parts using selective laser sintering (SLS) are the direct and indirect methods.

Fabricating parts with materials having high melting and sintering temperatures, especially in the case of bioceramics, has proven difficult not only because of thelimitations of the machines but also because of the difficulty of controlling the process parameters during fabrication [3, 4]. An alternative approach is an indirect SLS method, in which a binder either is used to fabricate a negative mold of the required porous structure or is mixed with the ceramic material to use as feedstock to the SLS machine [5-7]. The feasibility of fabricating scaffolds with 13-93 bioactive glass using an indirect SLS method with stearic acid (SA) as the binder was reported in our previous work [8]. In such indirect methods, not only should the binder be completely removed from the sintered part because of the biological application, but also the induced micro porosity should be eliminated as much as possible to improve the mechanical properties of the porous constructs.

Therefore, post-processing the fabricated “green parts” becomes crucial in such methods. In fact, post-processing steps also exist for a variety of additive manufacturing techniques such as stereolithography (SLA), 3D printing (3DP), fused deposition modeling (FDM), freeze extrusion fabrication (FEF), and others because of the inability of direct methods to process ceramics for similar reasons. Some of the important parameters that need to be considered for the fabrication of functional parts using such methods include the type of binder used, the actual material components and the thermolysis process [9]. The theory of de-binding has been researched for over two decades with reports on some qualitative and quantitative methods [10-13]. However, several aspects, including the wide range of particle sizes used and the complexity of the parts made using the SLS process, make it difficult to fit such parameters in the proposed quantitative research.

Therefore, we used an experimental approach to optimize the time required for heat treatment of “green” porous constructs in order to improve the densification when they are fabricated with different particle size ranges and binder content.In the current work, we investigate the effect of particle size, binder content, SLS parameters, de-binding and sintering heating rates, and immersion of scaffolds in simulated body fluid (SBF) on the microstructure of the scaffolds and on the compressive resistance of the sintered scaffolds. The materials used and methods of analysis are noted in Section 2. The results obtained, including the variation of compressive strengths with respect to the above-mentioned parameters, are presented in Section 3.2. Materials and Methods2.1.

The results obtained, including the variation of compressive strengths with respect to the above-mentioned parameters, are presented in Section 3.2. Materials and Methods2.1. Preparation of powder13-93 bioactive glass (prepared by Mo-Sci Corp., Rolla, MO) with a nominal chemical composition of 53% SiO2, 4% P2O5, 20% CaO, 5% MgO, 6% Na2O and 12% K2O (in weight %) was used in this research. The details of the 13-93 glass preparation are available elsewhere in the literature [4]. Two sets of 13-93 glass particles were prepared before mixing with the binder; the first set had an average particle size of ~44 µm, and the average particle size of the second set was ~0.7 µm. The first set of 13-93 glass particles were mixed with SA (C18H36O2, grade HS, Acros Organics, Morris Plains, NJ) and dry ball-milled for 8 hrs with ZrO2 grinding media to obtain the feedstock powder for the SLS machine containing 22% binder content by weight (hereafter represented as B22).

Then, a calculated amount of 13-93 glass from the first set was added to the B22 powder and dry ball-milled overnight to obtain the B18 powder. A similar procedure was used to prepare the B15 powder from B18 powder. The B12 powder was prepared by adding 13-93 glass, from the second set (with an average particle size ~0.7 µm), to the B15 powder and then dry ball-milling overnight. All the calculations were based on the bulk densitiesof the powders. Particle size distributions were obtained using a laser diffraction-based particle size analyzer (S3500, Microtrac Inc., Largo, FL). All the powders were examined by differential thermal analysis (DTA) (NETZSCH STA 409, Burlington, MA).2.2. Selective laser sinteringThe fabrication experiments were carried out on a commercial DTM Sinterstation 2000 machine.

Selective laser sinteringThe fabrication experiments were carried out on a commercial DTM Sinterstation 2000 machine. As mentioned earlier, an indirect SLS method was employed in this study, and a feasible set of SLS parameters conducive to fabricating scaffolds using 13-93/SA powders has already been established in our previous work [8]. All the parts in the current work were fabricated using a layer thickness of 76.2 µm. The part bed and part heater temperatures were maintained at 60°C, which is slightly below the melting point of SA (~69°C). This helps the SA to melt at low laser powers and avoids unnecessary temperature gradients in the part bed. The feed bin temperature was maintained at 35°C. An initial set of experiments was planned each time a powder with reduced binder content was used as feedstock to the SLS machine.

The feed bin temperature was maintained at 35°C. An initial set of experiments was planned each time a powder with reduced binder content was used as feedstock to the SLS machine. Several parts measuring 1 cm x 1 cm x 1 mm were initially fabricated to aid in determining the behavior of the powders, and the appropriate SLS control parameters, while being scanned by the laser beam. Our previous work established that an energy density of 1 cal/cm2 is required to consolidate successive layers. In this work, as powders with reduced binder amounts were used, the energy density was varied from 3.5 to 1 cal/cm2 at laser powers of 2W, 3W, 4W and 5W by maintaining the scan spacing at 228.6 µm. The 1 mm thick parts were visually inspected for any delamination, and the workable SLS parameter settings were determined based on the best feature definitions of the1 mm thick parts. The effects of SLS parameters will be discussed in greater detail in Section 3.4.2.3.

The effects of SLS parameters will be discussed in greater detail in Section 3.4.2.3. Post-processing and analysisThe fabricated green parts were post-processed in a three-stage programmable air furnace (Vulcan Benchtop Furnace, York, PA). Scanning electron microscopy (SEM) (S-570, Hitachi Co., Tokyo, Japan) images of the sintered parts were obtained to analyze the microstructures. A cross-head speed of 0.5 mm/min was used on a mechanical load frame (Instron 4469 UTM, Norwood, MA) to determine the compressive strengths of the parts. Five samples in each set were used, and the results are reported as the average value ± the standard deviation. Scans were run from 2θ values ranging from 10o to 80o using Cu Kα radiation (λ = 0.154056 nm) for powder X-ray diffraction (XRD) analysis (Philips X-Pert, Westborough, MA) on the sintered scaffolds and also on the dried scaffolds after removing them from the SBF to determine the crystalline/amorphous nature of the material.3.

Results and Discussion3.1. Powder characterization for the SLS processAs explained in Section 2.1, the particle size distribution was achieved by dry ball- milling the powders using ZrO2 grinding media. The graphical representation of the particle sizes is shown in Figure 1. The D50 particle size used in our previous work with B22 powders was ~42 µm. The B22 powder was dry ball-milled so that the D50 particle size decreased to ~24 µmin the current study; consequently, the layer thickness used in this study was reduced to 76 µm, from the 102 µm thickness used in our previous study. Figure 1 shows the distribution for the B18 powders, which is similar to a bimodal distribution with broad peaks at 10 µm and 44 µm. Improved compaction in the part bed would be obtained with this kind of distribution while spreading the powder in the SLS machine.

Improved compaction in the part bed would be obtained with this kind of distribution while spreading the powder in the SLS machine. It was expected in the preparation of the B15 powder that adding 13-93 glass particles to the B18 powder and dry ball-milling overnight would reduce the particle size. The peak in this case was shifted to 10 µm, however, with a significant amount of particles greater than 20 µm, which can be observed in its distribution shown in Figure 1. 13- 93 glass particles with an average particle size of ~0.7 µm were added to the B15 powder and then dry ball-milled overnight to obtain the B12 powder. The B12 particle size distribution showed 30% of the particles to be less than 5 µm and 70% of the particles to be less than 20 µm. The mean particle size values of the B18, B15 and B12 powders were 31 µm, 28 µm and 18 µm, respectively.

The mean particle size values of the B18, B15 and B12 powders were 31 µm, 28 µm and 18 µm, respectively. Therefore, it can be said that the mean particle size would not be a correct representation of the wide distribution of particle size ranges for the powders. This makes it difficult to include the particle size parameter in the quantitative approaches for the theory of de- binding because the powders behave differently during the heating cycles, as will be discussed later in the paper. The powders used to fabricate green parts are summarized in Table 1.Figure 1. Particle size distributions for powders with different binder contents.Table 1. Powder characterization Powders D50 (µm) D95 (µm) Mean (µm)B18 B15 B1220 16 1196 79 5631 28 183.2. Effect of heating ratesthe SLS parameters mentioned above, and using the B22 powder, are shown in Figure 2. Figure 2(a) shows the green parts, while the sintered parts are shown in Figure 2(b).

Effect of heating ratesthe SLS parameters mentioned above, and using the B22 powder, are shown in Figure 2. Figure 2(a) shows the green parts, while the sintered parts are shown in Figure 2(b). The parts fabricated include a designed porous part (60% porosity), a hollow part (wall thickness of 1.1 mm and 4The partsfabricated withFigure 2. (a) green parts (b) sintered partssymmetric holes measuring 1.2 mm in diameter) and a solid part with no designed porosity. All parts measured 6 mm in length and 5 mm in diameter. Hereafter, these parts will be referred to as porous, hollow and solid parts. Several batches of the hollow parts were fabricated to study the effect of heating rates on the microstructures, as well as on the compressive strength of the parts. The DTA/TGA analysis of the powders showed complete decomposition of the SA binder at ~540°C.

The DTA/TGA analysis of the powders showed complete decomposition of the SA binder at ~540°C. Therefore, the heat treatment of the green parts was completed in two phases: (i) a slower heating rate for de-binding the green parts from room temperature to 550°C and (ii) a relatively faster heating rate until reaching 700°C, with a1 hr hold for sintering of the glass particles followed by a furnace cool to room temperature. Two different sets of experiments were carried out to identify the best possible heating rates for both phases. In the first set of experiments, the sintering rate was kept constant at 2°C/min, and the de-binding rates varied from 0.1 to 2°C/min. In the second set of experiments, the effective de-binding heating rate of 1°C/min was chosen, and the sintering rate varied from 2 to 16°C/min. Figure 3 graphically represents the variation of the compressive strengths of the sintered hollow parts at different heating rates.

Figure 3 graphically represents the variation of the compressive strengths of the sintered hollow parts at different heating rates. The graphs indicate that the best possible heating rate before and after de-binding would be around 1°C/min and 5°C/min, respectively.Figure 3. Variation of compressive strength for different de-binding and sintering heating rates for the parts made with B22 powder.In all the indirect SLS techniques, controlling the heat treatment schedule during post- processing is an essential step. A slow heating rate increases the post-processing time, which could eventually decrease the attractiveness of the method. But, a fast heating rate during de- binding induces micro porosity because of the trapped gases, eventually producing flaws in the microstructure. From the theory of de-binding, it is generally understood that the strength of the parts decreases as the heating rate increases.

From the theory of de-binding, it is generally understood that the strength of the parts decreases as the heating rate increases. But, it can be observed from Figure 3 that the strength of the hollow parts increased up to a certain heating rate and then decreased. The faster de-binding heating rate is possible for the porous parts because of larger surface area available for the quicker oxidative decomposition of the SA binder. A further increase in the heating rate would rapidly melt the SA binder causing the porous part to collapse. Even if the part survives, there is the potential for bubble formation in the microstructures. Figure 4 shows the difference in microstructures for de-binding heating rates of 0.1, 1 and 1.5°C/min.

Even if the part survives, there is the potential for bubble formation in the microstructures. Figure 4 shows the difference in microstructures for de-binding heating rates of 0.1, 1 and 1.5°C/min. In our previous study a 3-day de-binding schedule was developed with holds and variable de-binding heating rates not higher than 0.2°C/min, and the microstructure obtained did not differ significantly from the microstructure obtained with a heating rate of 1°C/min (Figure 4b). Though the difference in compressive strength values is insignificant, this led to a greater reduction in the time requiredfor post processing. The effect of an increase in sintering heating rates on the compressive strength of the parts has a fairly narrow window of change, though it was observed that 4°C/min gave better results for the hollow parts, possibly because of their finer grain size [14].

A higher heating rate of 16°C/min caused deformation and slumping in the parts, which explains the large standard deviation. However, it must be noted that the heating rates are also dependent on the size of the parts. The developed heating rates were tested on a porous part measuring 1 cm in length and 1 cm in diameter. The applicability of the developed schedule on bigger parts must be investigated further and will be investigated in future studies.Figure 4. Effect of de-binding heating rates on microstructures (a-c) 0.1, 1 and 1.5°C/min.3.3. Effect of particle size and binder contentIt is known that a smaller particle size range results in a finer, sintered grain size and improved mechanical properties. However, smaller particles have larger surface areas and thus require higher amounts of binder content in the feedstock powder for the SLS machines, which could cause higher shrinkages after heat treatment.

Also, smaller particles with higher binder contents are known to cause spreading issues, resulting in warpage during the initial layers of fabrication and may require longer heat treatment schedules. The difficulty in using particle sizes less than 45 µm was reported by Goodridge et al. [16], which could be because of the larger layer thickness used (~ 250 µm), as increased thickness requires a high binder content and higher laser power to melt and join the successive layers. Therefore, a balance must be made between the particle size and the amount of the binder used during SLS fabrication. In our current work, the approach adopted was to decrease the binder content and the particle size simultaneously. In this way, workable parameters of particle size and binder content can be achieved with fewer experiments and less powder waste.

In this way, workable parameters of particle size and binder content can be achieved with fewer experiments and less powder waste. Green parts were successfully fabricated using B18 and B15 powders, whereas, the parts made with B12 were fragile to handle and difficulties were experienced while removing the unsintered powder from the pores. Although increasing the laser power to 5W improved the handling of the parts made with B12 powder, it still would not be feasible to handle bigger porous parts in different shapes. Moreover, these parts could not be sintered with the fast heat treatment schedule developed. Therefore, further reduction of binder content from the B12 powders was not considered as it would not be feasible.To study the variation of mechanical properties for reduced binder content and particle size, several porous, hollow and solid parts were fabricated with the powders prepared, as mentioned in Section 3.1.

The heat treatment schedule offering the best results, as discussed in Section 3.2, was used for binder burnout and for sintering the parts. Figure 5 showsrepresentative porous samples sintered at 700°C for 1 hr (heating at 1°C/min until 550°C and at 5°C/min until 700°C). The results show that the heat treatment schedule was successfully implemented for parts made with the B22, B18 and B15 powders but not with parts made with the B12 powder. According to the theory of thermal de-binding, for the case of oxidative degradation, inversely proportional to the particle size and packing densities [10]. As mentioned earlier, though the overall particle size range of all the powders is not significantly different, there was in increase in the number of particles that were less than 5 µm for the B12 powder, which gives it a better green packing density when compared to the other powders.

Therefore, though there is relatively little binder present in the B12 powder, because of the presence of smaller particle sizes, more time would be required for binder burnout; therefore, the accelerated heat treatment schedule did not work for this case. In contrast, higher green densities and wider particle distributions are often useful for better densification during the sintering process. A more appropriate heat treatment schedule for parts fabricated using the B12 powder must be developed with variable heating rates and hold times, which might be expected to take several days. Therefore, considering the poor handleability of these parts, combined with the unsuccessful heat treatment, they did not provide an optimal solution. Table 2 shows the variation of compressive strengths of the porous, hollow and solid parts made with the B22, B18 and B15 powders.

%) 60.1±24.7 22 76.9±22.5 18 115±28.4 15Porous (MPa) Hollow (MPa)18.3±5.8 19.4±5.4 22.1±3.840.9±6 43.7±12.2 45.9±6.33.4. Effect of SLS parametersBased on their higher overall compressive strengths, the B15 powder was used to study the effect of SLS parameters, again using the developed heat treatment schedule. In the powder- based, indirect AM techniques such as SLS and 3DP, the actual process is only used to fabricate a green part; it is understood that the strength of the sintered part depends on the heat treatment schedule. However, the presence of significant micro porosity in the green part could prove to be detrimental in achieving proper densification during the sintering process. Therefore, it is important to have a compact green part with little or no micro porosity unless the final goal is to achieve built-in porosity in the final sintered parts.

Therefore, it is important to have a compact green part with little or no micro porosity unless the final goal is to achieve built-in porosity in the final sintered parts. In our case, the porosity is designed in the CAD model, so the aim is to achieve maximum densification of the struts in the porous part to improve the mechanical properties of the final sintered part. The powders selected in this work, as explained in Section 3.1, were chosen to reduce the micro porosity in the green parts as much as possible.component of fabricating the green parts using the SLS process is the SLS process is the melting of the SA binder using a a laser beam, which not only has to fuse the successive layers but laser beam, which not only has to fuse the successive layers but the green parts.

This also has to fill in the voids between the particles to avoid micro porosity in the also has to fill in the voids between the particles to through changes in SLS process parameters such as laser power (P), beam speed can be achieved through changes in laser power (P), beam speed (BS) and scan spacing (SS). These three important parameters are connected using the following (BS) and scan spacing (SS). These three important parameters are connected using (BS) and scan spacing (SS). These three important parameters are connected using formula: Energy density, ED = P/(BS * SS) [15]. As the laser spot diameter for the DTM As the laser spot diameter for the DTM formula: Energy density, ED = P/(BS * SS) [ constant ~450 µm, scan spacing was kept constant at 228.6 µm, and the Sinterstation 2000 is a constant ~450 µm, scan spacing was kept constant at 228.6 µm constant ~450 µm, scan spacing was kept constant at 228.6 µm improved melting of other parameters were varied. The ant ED.

The ant ED. With the the SA binder: (i) increasing the lost by scanning the beam at low speeds while first method the sharpness of the features was lost by scanning the beam at low speeds first method the sharpness of the features This was believed to be due to melting of binder particles adjacent to the keeping P constant. This was believed to be due to particles adjacent to the part’s outline profile through heat transfer outline profile through heat transfer. This in turn affected the fabrication of designed the fabrication of designed not scanned by the laser would melt, thereby porous parts as the adjacent binder the parameters P and BS were decreasing the effective pore size minimum of 3W laser power was varied accordingly by keeping ED constant at 1 cal/cm laser power above 3W also required to avoid delamination of the fabricated parts. Increasing the laser power required to avoid delamination of the fabricated parts. maintaining the sharpness of the designed features.

Increasing the laser power required to avoid delamination of the fabricated parts. maintaining the sharpness of the designed features. Figure 6(a) melted the SA binder while maintaining he sharpness of the designed features. Figure 6(a) ows the fractured surface of the green part fabricated at 3W laser power using the B22 powder shows the fractured surface of the green part fabricated at 3W laser power using ows the fractured surface of the green part fabricated at 3W laser power using with the D50 particle size ~42 µm. Voids measuring up to ~200 µm in size are indicated in the are indicated in the with the D50 particle size ~42 µm. SEM image. In contrast, a fine distribution of the binder and partic le size with no presence of SEM image.

SEM image. In contrast, a fine distribution of the binder and partic le size with no presence of SEM image. In contrast, a fine distribution of the binder and particle size with no presence of such voids can be observed in Figure 6(b), which is the fracture surface of a part fabricated at such voids can be observed in Figure 6(b), which is the fracture surface of a part fabricated at such voids can be observed in Figure 6(b), which is the fracture surface of a part fabricated at increase in laser power could cause 5W laser power using the B15 powder. any voids present, the binder in the previously scanned la thereby reducing the micro porosity in the green part. cro porosity in the green part. Figure 6(c) shows the fracture surface of a Figure 6(c) shows the fracture surface of a sintered porous part fabricated at 5W laser power using the B15 powder. Similar observations Similar observations sintered porous part fabricated at 5W laser power using by Liu et al.

Similar observations Similar observations sintered porous part fabricated at 5W laser power using by Liu et al. [7], albeit using different materials and for dental were also made recently by Liu et al. albeit using different materials and for dental distribution of the binder in the green parts leads to a better applications. Therefore, a uniform oxidative degradation during the deThe other important component ofTherefore, there are two methods to achieve improved: (i) increasing the ED at constant P and (ii) increasing P at constant ED.binder particles not scanned by the laser wouldpore size in the green part. Therefore, the parameters keeping ED constant at 1 cal/cm2. A minimum of 3WB15 powder. This is because an increase in laser power couldin the previously scanned layers to re-melt and distribute itself into anyTherefore, a uniform distribution of the binder in the green parts lead ng densification.

during the de-binding stage, thereby improving densificationFigure 6. SEM images of the fractured surfaces (a) green part made with B22 powder (b) green SEM images of the fractured surfaces (a) green part made with B22 powder (b) green SEM images of the fractured surfaces (a) green part made with B22 powder (b) green part made with B15 powder (c) B15 sintered part showing improved densification. part made with B15 powder (c) B15 sintered part showing improved densification. part made with B15 powder (c) B15 sintered part showing improved densification.Figure 7 shows the improvements in the mechanical properties of the porous, hollow and in the mechanical properties of the porous, hollow and Figure 7 shows the improvement at a constant energy density of 1 cal/cm2 with varying laser powers of 3W, with varying laser powers of 3W, beam speeds of 305, 406 and 508 mm/s. All the parts were made 305, 406 and 508 mm/s. All the parts were made schedule developed in Section 3.2.

All the parts were made 305, 406 and 508 mm/s. All the parts were made schedule developed in Section 3.2. The MPa to 41 MPa, was increased from 3W to 5W. Similar trends can also be seen for the Similar trends can also be seen for the MPa and 157 MPa,solid parts fabricated at a constant energy density of 1 cal/cm 4W and 5W and corresponding beam speeds of using the B15 powder and were compressive strength of the porous parts almost doubled, increasing from 22 compressive strength of the porous parts almost when the laser power was increased hollow and solid parts with thewere sintered using the schedule developed in Section 3.2he compressive strengths increasing to 61.4 MParespectively. The compression modulus values for the porous, hollow and solid parts were 4.4±0.7 GPa, 6.1±1.6 GPa and 12±1 GPa, respectively.

The compression modulus values for the porous, hollow and solid parts were 4.4±0.7 GPa, 6.1±1.6 GPa and 12±1 GPa, respectively. The modulus values of the porous and hollow parts are similar to that of human trabecular bone and values of solid parts are similar to that of cortical bone [16].Figure 7. Effect of laser power and beam speed on compressive strength of partsfabricated at a constant energy density of 1 cal/cm2.3.5. SBF testsThe SBF solution was prepared based on the Kokubo method [17]. All the samples were ultrasonically cleaned three times using distilled water and then ethanol. The samples were dried in an oven maintained at 65°C overnight and then weighed before being kept in the SBF solution. 100 ml of solution was used for 1 mg of the scaffold. The SBF solution container with scaffolds was kept in an incubator maintained at 37°C.

100 ml of solution was used for 1 mg of the scaffold. The SBF solution container with scaffolds was kept in an incubator maintained at 37°C. All the compression tests were conducted on wet scaffolds to provide realistic data on the degradation of the scaffolds over different time periods. The variation of the compressive strengths from 1 to 6 weeks is provided in Figure 8. A detailed discussion of the behavior of 13-93 glass and its degradation in SBF can be foundFigure 8. Variation of compressive strength of the samples immersed in SBF.study the variation of elsewhere in the literature [18]. The main objective of this method after immersion in mechanical properties of the 13- the following heat SBF.

The main objective of this method after immersion in mechanical properties of the 13- the following heat SBF. The scaffolds used for this study were made using treatment schedule: de-binding heating rate of 1°C/min until 550°C and then heating rate of binding heating rate of 1°C/min until 550°C and then heating rate of binding heating rate of 1°C/min until 550°C and then heating rate of ntering temperature of 700°C was reached. The compressive strength of the 2°C/min until the sintering temperature of compressive strength of the scaffolds measured after 6 weeks was 11 MPa, approximately 38% reduction, and is at porous scaffolds measured after 6 weeks was 11 MPa, approximately 38% reduction, and is scaffolds measured after 6 weeks was 11 MPa, approximately 38% reduction, and is the higher end of the compressive strength of a human trabecular bone [19]. the higher end of the compressive stren.

the higher end of the compressive stren. The main objective of this work was to study -93 glass scaffolds made using SLS method after he scaffolds used for this study were made using the B22 powder and the, sputter coated with gold- After testing the scaffolds the scaffolds were dried at room temperature, sputter coated with gold using SEM. Specifically, SEM was used to look for HA formation for HA formation. palladium, and investigated using The HA crystals were formed within a week of the scaffolds’ immersion in SBF, which is in immersion in SBF, which is in The HA crystals were formed within a week of the Figure 9 shows the accordance with other reports of the in vitro behavior of 13 SBF after 2 and 6 weeks, SEM images of two representative respectively.

Figure 9 depicts the typical rough surface morphology of the SLS scaffolds (seen in he typical rough surface morphology of the SLS scaffolds (seen in Figure 6(c)) converting to a smooth surface with a crystalline layer within 2 weeks of immersion. Figure 6(c)) converting to a smooth surface with a crystalline layer within 2 weeks of immersion. Figure 6(c)) converting to a smooth surface with a crystalline layer within 2 weeks of immersion. Figure 9(b) shows a fractured surface of a strut of the porous sample after immersion in SBF for Figure 9(b) shows a fractured surface of a strut of the porous sample after immersion in SBF for Figure 9(b) shows a fractured surface of a strut of the porous sample after immersion in SBF for distinguished on the surface of the 13- formed by crystals, can be clearly distinguished on the surface of the 13 6 weeks. A thick layer, formed by SEM image shown in Figure 9(c).

A thick layer, formed by SEM image shown in Figure 9(c). XRD analysis run on the sample 93 glass in the magnified SEM XRD analysis run on the sample typical amorphous hump of the 13-93 glass and showed peaks similar to showed a loss of the typical amorphous hump of the 13 93 glass and showed peaks similar to that of HA [4]. The growth and differentiation of osteoblasts on 13 The growth and differentiation of osteoblasts on 13-93 scaffolds prepared using 93 scaffolds prepared using other, traditional methods such as polymer foam replication has already been reported [20-22]. other, traditional methods such as polymer foam replication has already been reported [20 other, traditional methods such as polymer foam replication has already been reported [20 93 scaffolds made using SLS process offer vigorous Our previous study also showed that 13-93 scaffolds made using SLS process of Our previous study also showed that 13 growth of MLO-A5 cells.

With improved mechanical properties and SBF test results, combined A5 cells. With improved mechanical properties and SBF test results, combined A5 cells. With improved mechanical properties and SBF test results, combined with the flexibility of manufacturing anatomical shapes as described in the present paper, the with the flexibility of manufacturing anatomical shapes as described in the present paper, the with the flexibility of manufacturing anatomical shapes as described in the present paper, the indirect SLS process shows its high potential for manufac turing highly interconnected porous indirect SLS process shows its high potential for manufacturing highly interconnected porous structures and load bearing scaffolds for bone repair applications using bioactive glasses. structures and load bearing scaffolds for bone repair applications using bioactive glasses.

structures and load bearing scaffolds for bone repair applications using bioactive glasses. structures and load bearing scaffolds for bone repair applications using bioactive glasses.other reports of the in vitro behavior of 13-93 glass [20, 21]. Figuretwo representative scaffolds taken out of the SBF after 2Figure 9. (a) Surface morphology of a porous sample after immersion in SBF for 2 weeks (b) urface morphology of a porous sample after immersion in SBF for 2 weeks (b) urface morphology of a porous sample after immersion in SBF for 2 weeks (b) Fracture surface of a porous sample after immersion in SBF for 6 weeks (c) Magnified image sample after immersion in SBF for 6 weeks (c) Magnified image sample after immersion in SBF for 6 weeks (c) Magnified image showing the growth of HA film on the surface. showing the growth of HA4.

showing the growth of HA4. Summary and ConclusionThe main research objective of this work was to study the effects of material, thermal and effects of material, thermal and The main research objective of this work was to study the SLS process parameters on the mechanical properties of 13-93 bioactive scaffolds manufactured 93 bioactive scaffolds manufactured SLS process parameters on the m using indirect selective laser sintering (SLS) process. The primary aim of this work was to primary aim of this work was to using indirect selective laser sintering (SLS) process 93 bioactive scaffolds. This was achieved by reducing improve the mechanical properties of 13-93 bioactive scaffolds.

This was achieved by reducing improve the mechanical properties of 13-93 bioactive scaffolds. This was achieved by improve the mechanical properties of 13 binder content of the feedstock powder, varying the SLS process parameters the particle size and binder content of the feedstock powder, varying the SLS process parametersduring fabrication and improving the heat treatment schedule of the green parts. The approach used in this work reduced the particle size and the amount of binder content simultaneously. The results showed that the minimum amount of stearic acid binder required to mix with 13-93 glass to obtain a workable feedstock powder for SLS machine was 15% by weight (B15 powder). The particle size distribution with a D50 size of ~16 µm and D95 size of ~79 µm obtained for the B15 powder gave the best possible fabrication results with reduced micro porosity in the green part.

The particle size distribution with a D50 size of ~16 µm and D95 size of ~79 µm obtained for the B15 powder gave the best possible fabrication results with reduced micro porosity in the green part. The heat treatment schedule developed with a de-binding heating rate of 1°C/min and a sintering heating rate of 5°C/min was based on the highest compressive strengths achieved for the set of experiments as explained in Section 3.2. Although there was an insignificant increase in the mechanical properties, the time required for post processing the green parts was significantly reduced from 80 hrs to 12 hrs. Using the developed heat treatment schedule, the highest compressive strength for the porous sample geometry was achieved for the parts made with B15 powder. Therefore, the B15 powder was used to study the effect of SLS process parameters.

Therefore, the B15 powder was used to study the effect of SLS process parameters. It was observed that at a constant energy density of 1 cal/cm2, increasing the laser power and beam speed from 3W and 305 mm/s to 5W and 508 mm/s had a significant effect on the mechanical properties of the parts. The compressive strength of the porous parts made with B15 powder was further increased to 41 MPa from 22 MPa. However, because of time constraints, SBF tests were carried on the parts made with B22 powder using non-optimized parameters, i.e., fabricated at 3W laser power; 305 mm/s beam speed; 1 cal/cm2 energy density and post-processed at 1°C/min de-binding heating rate and 2°C/min sintering heating rate. The compressive strengths of these dry porous parts are indicated in the trend of compressive strength vs. soaking time shown in Section 3.5 (see Figure 8).

The compressive strengths of these dry porous parts are indicated in the trend of compressive strength vs. soaking time shown in Section 3.5 (see Figure 8). Assuming a similar trend, it can be estimated that the scaffolds with improved mechanical properties (~41 MPa for ~60% porosity) would have a compressive strength ~25 MPa after 6 weeks, which is much higher than the human trabecular bone.It is concluded that the presence of sub-micron sized particles in the feedstock powder not only increases the post processing time to several days but also requires higher binder content during fabrication. Increasing the laser power and beam speed at a constant energy density during SLS fabrication not only improves the densification in the green part but also preserves the sharpness of the features fabricated when compared to increasing the energy density at a constant laser power and low beam speeds.

The compressive strengths of the parts fabricated with the reduced binder content and particle size, improvised heat treatment schedule and SLS parameters varied from 41 MPa for a scaffold with ~60% porosity to 157 MPa for a solid part with no designed porosity. The improved mechanical properties combined with the promising in vitro results show the high potential of scaffolds manufactured with 13-93 bioactive glass using the selective laser sintering process in load bearing applications for bone repair.AcknowledgementsThis research was funded by the U.S. Navy under SBIR phase I award N00014-11-M- 0113 to Mo-Sci Corporation and by the Intelligent Systems Center at Missouri University of Science and Technology. The authors thankfully acknowledge the help of Xin Liu and Kevin Wu during various stages of this work.References[1] Hench L. L. and Wilson J., “An Introduction to Bioceramics” (Singapore: World ScientificPublishing Co. Pte. Ltd.) p 47 (1993).[2] Rahaman M.

fective Cooling Method for Spin Casting ProcessYong-Ak Song, Sehyung Park, Yongsin KwonKorea Institute of Science and Technology KIST, CAD/CAM Research Center P.O. Box 131, Cheongryang, Seoul, KoreaAbstractSpin casting has been widely used in prototyping industry as a secondary process to convert a master model into a functional metal or plastic part. The main problem of the spin casting process consists in the poor thermal conductivity of silicone rubber as mold material which leads to a long cooling time between each casting processes and also to a short life time of mold. To solve this problem, different cooling methods have been developed and compared to each other experimentally. First, air cooling channel has been integrated into the spin casting mold to enhance the heat removal. Secondly, the silicone rubber has been mixed with different metal and ceramic powders to increase its thermal conductivity.

Secondly, the silicone rubber has been mixed with different metal and ceramic powders to increase its thermal conductivity. The results so far prove an applicability of the developed cooling methods.1. IntroductionSpin casting is a secondary process to produce metal and plastic parts quickly from an existing master model. It is especially suitable for the manufacture of die casting parts since the manufacture of dies is normally very time-consuming and expensive. In spin casting, a silicone rubber material is used. Since the silicone rubber is a flexible material before vulcanizing, it can be easily formed around the master model, thus making the mold manufacture fast and easy to learn.Although the silicone rubber material can withstand a temperature of 420 °C and thus allowing the use of low-melting alloys such as tin and zinc for casting, it has a low thermal conductivity. A typical value for the thermal conductivity of rubber is _=0.2 W/Km [1].

A typical value for the thermal conductivity of rubber is _=0.2 W/Km [1]. This low thermal conductivity leads to an increasing mold temperature in the course of casting process. The constant heating-up of the mold has a damaging effect on the mold surface since it can be worn off due to the thermal stress or even burned off. To avoid such a damage, it is necessary to cool down the mold surface between each shots with compressed air or just with natural convection. This, however, increases the total cycle time. Also, according to [2], the deviation increases with higher mold temperature. For this reason, it is necessary to cool down the mold during spin casting, thereby keeping the mold temperature constant.552. Development of cooling methodsIn the experiment, a spin casting equipment of Tekcast is used.

Development of cooling methodsIn the experiment, a spin casting equipment of Tekcast is used. Since the mold rotates during the casting process, it is rather difficult to supply a constant flow of coolant from the outside into the mold as it is mostly done in conventional injection molding. A closed loop of coolant flow also imposes a serious problem regarding the proper sealing of the mold. Under the circumstances, air seems preferable to water. Also in view of easier integration into the existing spin casting equipment, air cooling seems a better choice than water cooling. To realize an air cooling system, the center of the shaft coaxial to the pneumatic cylinder is modified to let the air flow into the inside of the mold, as shown in Fig. 1. To connect the lower and the upper mold, a plastic tube is used. By this means, the air flows from the lower part of the mold to the upper part.Fig.

1. To connect the lower and the upper mold, a plastic tube is used. By this means, the air flows from the lower part of the mold to the upper part.Fig. 1: Flow of cooling air through the moldThe way how the mold is fabricated offers a great flexibility to design the cooling channel into the inside of the mold. According to the thickness of a master model used, an appropriate number of silicone rubber discs are stacked on each other. To implement an internal cooling channel in the mold, the copper pipes are then inserted between those silicone rubber discs, preferably directly beneath the cavity surface, Fig. 2. The copper pipes can be bended prior to the insertion and in this way different shapes of the cooling channel can be realized. After vulcanizing of the silicone rubber discs, the pipes are then firmly included in the mold.

After vulcanizing of the silicone rubber discs, the pipes are then firmly included in the mold. With regard to a high thermal conductivity around the cooling channel, it is recommendable to let the copper pipes in the inside of the mold.56In order to enhance the cooling effect in the mold during air cooling, it is tried to increase the thermal conductivity of the mold locally around the cavity. A commonly used method for this purpose is to add a high thermal conductivity material to the mold material. For example, a high thermal conductivity metal powder such as aluminum is added to the UV-curable resin which is then used in SLA process to make a mold for low temperature molding process [3, 4]. Two metal powders, copper (_=350 W/Km) and aluminum(_=221 W/Km), as well as two ceramic powders, alumina (_=27.6 W/Km) and silicon nitride (16-33 W/km), are used in the experiment.

Two metal powders, copper (_=350 W/Km) and aluminum(_=221 W/Km), as well as two ceramic powders, alumina (_=27.6 W/Km) and silicon nitride (16-33 W/km), are used in the experiment. Although ceramics have lower thermal conductivities compared to those of metals, ceramics offer a higher heat resistance, thus protecting the mold surface better than metals at higher casting temperature. The powder size of used powder materials is between 15-40 _m.Fig. 2: Copper-mixed cavity and cooling pipes in the spin casting mold3. Comparison of different cooling methodsTo measure how effective each cooling methods are, molds with and without internal cooling channel are compared to each other regarding their surface temperatures. The experimental parameters are as follows:Clamping pressure: Rotating speed: Melting temperature of zinc: Pouring time: Flow rate of cooling air:4000 PSI 500 rpm 410 ºC 1 min.

The experimental parameters are as follows:Clamping pressure: Rotating speed: Melting temperature of zinc: Pouring time: Flow rate of cooling air:4000 PSI 500 rpm 410 ºC 1 min. 140-150 l/minIn case of in-process cooled molds, the cooling effect of the combination of air cooling with the copper-mixed cavity is compared to that of the solely air cooled mold as well as to that of the copper-mixed cavity without air cooling. The weight mix ratio of silicone rubber to copper is 3:2. After each shot, the surface temperature of the molds is measured. As shown in Fig. 3,57the surface temperature of a mold without any cooling reaches 140 °C after 20 shots and then remains almost constant at that temperature. If air is supplied to the mold, the surface temperature is reduced to 110 °C.

If air is supplied to the mold, the surface temperature is reduced to 110 °C. This result means that even a simple air supply into the mold is an effective cooling method.1 6 0]C1 5 0Cu-mixed cavity[1 4 0e r u t a r e p m e te c a f r u s1 3 01 2 01 1 01 0 0silicone rubbersilicone rubber + air coolingd l o M9 08 0Cu-mixed cavity + air cooling7 01 2345 67 8 9 1 0 1 1 1 2 1 3 1 4 1 5 1 6 1 7 1 8 1 9 2 0Number of shotsFig.3: Surface temperatures of molds cooled with different cooling methodsThe high surface temperature of the copper-mixed mold around 160 °C proves that the addition of copper to silicone rubber increases the thermal conductivity of the cavity and thus making it a heat sink. When cooled with air, the surface temperature drops to 100 °C, which is around 10 °C lower than that of the pure silicone rubber mold cooled with air.

When cooled with air, the surface temperature drops to 100 °C, which is around 10 °C lower than that of the pure silicone rubber mold cooled with air. This result implies that the combination of copper-mixed mold with air cooling can be used if air cooling of silicone rubber mold alone is not sufficient enough.In case of aluminum-mixed molds, the surface temperature also increases to around 160 °C if no air cooling is additionally used, Fig. 4. When air is supplied, the surface temperature decreases to around 120 °C. Compared to that of the pure silicone rubber mold, however, this is no distinct temperature difference, the reason being that the mix ratio of silicone rubber and aluminum is not optimal yet. The ceramic-mixed molds also show no distinct cooling effect compared to that of the pure silicone rubber mold. However, the ceramic-added molds show less surface burn after several shots.

The ceramic-mixed molds also show no distinct cooling effect compared to that of the pure silicone rubber mold. However, the ceramic-added molds show less surface burn after several shots. Therefore, in terms of the surface protection, the addition of ceramic powder gives a distinct advantage over the use of pure silicone rubber mold.581 7 01 6 0]C[e r u t a r e p m e te c a f r u s1 5 01 4 01 3 01 2 0alumina-mixed mold air coolingalumina-mixed moldaluminum-mixed mo + air coolingaluminum-mixed mod l o M1 1 0silicon nitride + air cooling1 0 09 08 01 2 34 5 67 8 9 1 0 1 1 1 2 1 3 1 4 1 5 1 6 1 7 1 8 1 9 2 0Number of shotsFig. 4: Surface temperatures of molds mixed with different materials4. ConclusionIt was aimed to develop an in-process cooling method for spin casting process.

4: Surface temperatures of molds mixed with different materials4. ConclusionIt was aimed to develop an in-process cooling method for spin casting process. Two cooling methods are proposed: First, an insertion of internal cooling channel with copper pipe for air cooling and second, an increase of thermal conductivity of mold by mixing the silicone rubber with a high thermal conductivity material. The result of the investigation shows that a cooling effect can be already achieved when compressed air is supplied into the mold through the copper cooling pipes. By this means, the surface temperature can be decreased by around 30 °C. The cooling effect can be enhanced if materials of higher thermal conductivity such as copper are added to the silicone rubber in the cavity area. Among the tested materials, copper shows the best cooling effect. In case of the copper-mixed mold with air cooling, the surface temperature is even reduced by around 40 °C.

Among the tested materials, copper shows the best cooling effect. In case of the copper-mixed mold with air cooling, the surface temperature is even reduced by around 40 °C. This result clearly indicates the effectiveness of the developed cooling methods. Regarding the surface protection during spin casting, ceramic materials offer a distinctive advantage over metals and pure silicone rubber mold. In future work, the influence of mold cooling temperature on the part accuracy has to be investigated in detail.595. References[1] http://www.mininglife.com/Miner/general/thermalconductivity.htm[2] Barnard, L.J., “Spin casting as a tool in rapid prototyping”, Solid Freeform Fabrication 1999, The University of Texas at Austin, 9-11 August 1999, pp. 719-725[3] Girouard, D., “Molds for low temperature molding processes using rapid prototyping”, The Fourth International Conference in Rapid Prototyping, 14-17 June 1993, Dayton, Ohio, p.

719-725[3] Girouard, D., “Molds for low temperature molding processes using rapid prototyping”, The Fourth International Conference in Rapid Prototyping, 14-17 June 1993, Dayton, Ohio, p. 61-66[4] Nakagawa, T., “Rapid Prototyping in Japan”, The Eighth International Conference on Rapid Prototyping, 12-13 June 2000, Tokyo, pp. 1-1660.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEffective Elastic Properties of Additively Manufactured Metallic Lattice Structures: Unit-cell ModelingO. Fashanu, D. Murphy, M. Spratt, J. Newkirk and K. Chandrashekhara Missouri University of Science and Technology, Rolla, MO 65409B. Brown and J. Porter Department of Energy’s Kansas City National Security Campus, Kansas City, MO 64147AbstractLattice structures are lightweight materials, which exhibit a unique combination of properties such as air and water permeability, energy and acoustic absorption, low thermal conductivity, and electrical insulation. In this work, unit-cell homogenization was used to predict the effective elastic moduli of octet-truss (OT) lattice structures manufactured using selective laser melting (SLM).

In this work, unit-cell homogenization was used to predict the effective elastic moduli of octet-truss (OT) lattice structures manufactured using selective laser melting (SLM). OT structures were manufactured using a Renishaw AM 250 SLM machine with various relative densities. Compression test was carried out at strain rate 5 × 10-3 m-1 using an MTS frame. Finite element analysis was used in the determination of the OT’s effective elastic properties. Results from the finite element analysis were validated using experiments. It was observed that the finite element predictions were in good agreement with the experimental results. This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing & Technologies, LLC under contract number DE- NA0002839.1.

IntroductionThe development and continuous improvement of additive manufacturing methods have dramatically changed the way designers and engineers pursue design and manufacturing [1]. Additive manufacturing, widely known as 3D-printing, is a method that forms parts from powder, wire, or sheets in a process that proceeds layer-by-layer [2]. This process has been used in many fields for concept modeling; for example, to manufacture customized and lightweight cellular structures.Cellular structures, which include honeycombs, lattice structures, and foams, are lightweight structures that exhibit excellent mechanical, thermal, and acoustic properties. These structures are excellent candidates for applications, which require a high strength to weight ratio as used in the aerospace industry [3-4].

These structures are excellent candidates for applications, which require a high strength to weight ratio as used in the aerospace industry [3-4]. The properties of cellular structures are dependent on three factors: the topology of the cellular structure, the properties of the parent material, and the relative density ((cid:591)(cid:3364)) of the lattice structure. The mechanical performance of cellular structures are usually estimated by the effective properties. Considering the same material and relative density, the topology controls the deformation mode (bending or stretching) of the structure. According to Gibson and Ashby [5], the strength of bending dominated structures for all loading conditions scales with (cid:591)(cid:3364)1.5, while the strength of stretching dominated structures scale with (cid:591)(cid:3364). For example, for a relative density of 0.2, a stretching dominated structure is expected to be about two times as strong as a bending dominated structure.

For example, for a relative density of 0.2, a stretching dominated structure is expected to be about two times as strong as a bending dominated structure. The topology space (i.e. different types of topologies)2223has not been fully investigated due to the cost of trial and error experimentations. Numerical homogenization, a less expensive alternative is proposed in this study as a possible way of investigating different topologies.In this study, a unit-cell homogenization model was developed to predict the effective elastic properties of an octet-truss structure. The unit-cell homogenization predictions were compared against experimental results to review the accuracy of the homogenization method.2. TheoryNumerical homogenization has been used to predict the mechanical properties of composites [6-7]. It involves the transformation of a heterogeneous solid into a homogenous solid.

TheoryNumerical homogenization has been used to predict the mechanical properties of composites [6-7]. It involves the transformation of a heterogeneous solid into a homogenous solid. The effective elastic properties of the homogenous body is determined by the relationship between the average stress and average strain in a unit-cell or representative volume element (Eq. 1).

Methodology3.1. Manufacturing of test specimensA Renishaw AM250 SLM machine was used to manufacture two different octet-truss (OT) structures. OT-A was designed with a relative density ((cid:591)(cid:3364)) of 0.05 (Fig. 1a) and cell size of 4.56 mm, while OT-B was designed with (cid:591)(cid:3364) of 0.2 (Fig. 1b) and cell size 4.56 mm. Both OT structures had a configuration of 6 × 6 × 10 unit cells (Fig. 2).2224(1)(2)(3)(a)(b)Figure 1. OT unit-cell (a) relative density 0.05, (b) relative density 0.20(a)(b)Figure 2. Manufactured OT structures (a) relative density 0.05, (b) relative density 0.203.2. Mechanical testingMechanical compression testing of the manufactured OT structures was performed using an MTS 380 frame. The crosshead speed was set at 0.228 mm/min, resulting in a nominal strain rate of 5 × 10-3min-1. Three specimens were tested for each OT structure type.

The crosshead speed was set at 0.228 mm/min, resulting in a nominal strain rate of 5 × 10-3min-1. Three specimens were tested for each OT structure type. After the compression testing, the effective elastic moduli were estimated from the resulting true stress-strain curves.4. Results and Discussion4.1. Experimental resultsAt the completion of each compression test, the true stress-strain curves were obtained by dividing the force by the actual cross-sectional area. The actual cross-sectional area of the OT structure was assumed to increase in the same manner as a solid under compression. Fig. 3 shows the true stress- strain curves of OT-A ((cid:591)(cid:3364) = 0.05) and OT-B ((cid:591)(cid:3364) = 0.20). It can be observed that the stress-strain curve of OT-A is different in terms of behavior from OT-B. The difference in behavior can be attributed to the difference in strut thicknesses.

It can be observed that the stress-strain curve of OT-A is different in terms of behavior from OT-B. The difference in behavior can be attributed to the difference in strut thicknesses. The thin struts of OT-A allow it to deform differently when compared to OT-B which had thicker struts. Fig. 3 also shows that the strength and effective modulus increases as the strut thickness increases. The mechanical properties of the two different OT structures are presented in Table 1.Table 1. Mechanical properties of SLM 304L stainless steel OT structuresOT-AOT-B2225Relative density ((cid:2761)(cid:3365)) Effective strength (MPa) Effective modulus (GPa)0.05 4.28 ± 0.1 0.61 ± 0.030.20 34.28 ± 0.2 2.77 ± 0.01660550~4 ~ ~3~40 ~ ~30[/J <I).b 2 VJ[/J <I).b 20 VJ10000.20.40.60.8000.20.40.60.8Strain [mm/mm] (a)Strain [mm/mm] (b) Figure 3. True stress-strain curve (a) OT-A, (b) OT-B4.2.

True stress-strain curve (a) OT-A, (b) OT-B4.2. Numerical homogenizationIn this study, the unit-cell homogenization method was used to predict the effective elastic properties of the OT structure (Fig. 1). The unit-cell was selected because of the low computational cost associated with it. The OT structure consists of an octahedral cell as well as eight tetrahedral cells (Fig. 4). The heterogeneous body was assumed to consist of the combination of the octahedral cell and air. The octahedral cell alone was considered for this analysis because this cell absorbs all the loads within the OT structure [8-9].(b) Figure 4. (a) Octet-truss structure, (b) octahedral cell, (c) tetrahedral cell(a)(c)Numerical homogenization of the OT unit-cell was carried out using ABAQUS commercial software. The used input parameters are shown in Table 2. The mesh used consisted of approximately 94948 tetrahedron solid elements with an average global size of 0.2, as shown in Fig. 5.2226Figure 5.

The effective elastic modulus for the two different OT structures were evaluated using the relationship between the stiffness matrix, compliance matrix, and elastic constants. Fig. 6 and Table 3 show the comparison between the predicted unit- cell homogenization results and the experimental results.~3.5 0-. d3'---' rJ) ;:::l '3 2.5 '"O 0 s2 ◊I ◊ Experiment * Homogenization model! I0.050.10.150.20.25Relative densityFigure 6. Comparison between the numerical model and experimentFrom the figure, it can be observed in both cases (numerical model and experiment) that an increase in relative density led to an increase in stiffness. This increase in stiffness can be attributed to the increase in strut thickness. Increasing the relative density led to an increase in strut thickness,2227and, a larger strut is expected to have a higher stiffness.

Increasing the relative density led to an increase in strut thickness,2227and, a larger strut is expected to have a higher stiffness. From Table 3, it can also be observed that the elastic moduli predicted by the unit-cell homogenization model was within 10% of the experimental results. Therefore, it can be concluded the unit-cell homogenization model can be used to predict the effective elastic properties of cellular structures.Table 3. Comparison between numerical model and experimentPropertyHomogenization modelExperimentRelative density0.050.200.050.20E*11 (GPa)0.563.060.612.775. ConclusionIn order to numerically predict the effective elastic properties of lattice structures manufactured using SLM, a unit-cell homogenization approach was investigated. Two octet-truss structures with different relative densities (0.05 and 0.20) were manufactured and tested.

Two octet-truss structures with different relative densities (0.05 and 0.20) were manufactured and tested. The unit-cell homogenization model was used to predict the effective stiffness of these two structures. From the results, it was observed that the numerical unit-cell predictions were within 10% of the experimental results. As such, it can be concluded that the unit-cell homogenization approach can be used to effectively predict the effective stiffness of lattice structures.6. AcknowledgmentsThis work was funded by the Department of Energy’s Kansas City National Security Campus, which is operated and managed by Honeywell Federal Manufacturing & Technologies, LLC under contract number DE-NA0002839.7. References[1]N. K. Rajak and A. Kaimkuriya, “Design and Development of Honeycomb Structures for Additive Manufacturing,” International journal of Trend in Scientific Research and Development (IJTSRD), vol. 2, pp. 1198–1203, 2018.[2] W. J. Sames, F. A. List, S. Pannala, R.

fective MechanisUls of Multiple LED Photographic CuringKai Loose, Toshiki Niino The Institute of Physical and Chemical Research (RIKEN)Takeo Nakagawa Institute of Industrial Science, University of TokyoAbstractMultiple LED Photographic Curing (MPC) has proven capable of drawing cross sections of three-dimensional objects like printing a sheet of paper. Using raster scanning, however, simulta neously exposing a photopolymer with 1,024 beams of light involves various unknown issues. The aim of this research work was to examine the formation of individual strings and the connecting mechanisms between strings and layers. At light power ranging from 19.0 to 30.3flW and at vari ous scan speeds, string formation perpendicular to scan direction differs greatly from that in scan direction.

At light power ranging from 19.0 to 30.3flW and at vari ous scan speeds, string formation perpendicular to scan direction differs greatly from that in scan direction. Curing of plane layers happens by curing strings side-by-side with a constant spacing of 62.5flm.1IntroductionMultiple LED photographic curing (MPC) is based on the concept of exposing a liquidphotopolymer through multiple beams of light simultaneously. Unlike Laser Stereolithography (SL), MPC uses light-emitting diodes (LED) as light sources (Loose et aI., 1997). Exposure hap pens through an array of 1,024 beams of light which expose the surface of the liquid resin while using raster scanning. Similar concepts for multiple-spot exposure have been proposed byOht subo et al. (1997) and Marutani et al. (1998). All concepts have successfully shown their practical applicability.Nevertheless, very little is known about the effective mechanisnls in multi-spot exposure.

(1997) and Marutani et al. (1998). All concepts have successfully shown their practical applicability.Nevertheless, very little is known about the effective mechanisnls in multi-spot exposure. In spite of seemingly principal compatibility between MPC and SL-both processes are essentially an exposure of a liquid photosensitive resin leading to polymerisation, thus to solidobjects-··there is a number of challenges to be met considering the unique characteristics of the exposure pattern. It can be said that the research on MPC is at the developmental stage, and large.number of nleas urements are required to improve the design. For example, the effective mechanisms when ex posing the photopolymer resin surface are unknown. If they are understood to some extent, MPC can be improved to become applicable· for rapid concept Inodelling.This paper describes the results of the experiluental examination of the curing process in MPC.

Not only the curing of single strings but also the formation of layers and three-dimensional177objects is studied. Consequently, the authors propose a means to generate an equal cure depth through multiple beams of light and investigate it theoretically and experimentally. In the follo~­ ing, experimental results are presented and the envisaged application explained. The results of this research indicate thatMPC is a successful method for faithfully generating three-dimensional ob jects, and has the potential to aid design evaluation.2 Background of the ExaminationsToday, there is a growing demand for cheap and fast concept modelling. Research on MPCcarried out at RIKEN aims at combining this application with the unique characteristics of Stereolithography parts, such as transparency.

Research on MPCcarried out at RIKEN aims at combining this application with the unique characteristics of Stereolithography parts, such as transparency. Using inexpensive technology, MPC is an attrac tive technique to process photopolymers into various shapes.In the frame of the following examination, the decisive process-technological parameters ofinfluence shall be investigated. Figure 1 explains that the following examinations focus on the features which are characteristic of the MPC process.

Together with an adjusted control software, the way how energy is absorbed in the resin surface is of utmost importance regarding the achiev able accuracy to form and shape.Energy supply - Multi-spot exposure - Multimode intensity distributionSubjects - Forming of strings - Connecting mechanisms betweenin a single beam of lightstringsPresence of multiple beams of light simultaneously exposing the photopolymerControl software - Selective curing of a bitmap image representing a part's cross-sectionConnecting mechanisms between layersInfluence of exposure time on process quality Influence of light power on process qualityExposure pattern is controlled byBuilding a part layer-by-Iayer - Selective curing of resin bymultiple-dot exposure using visible light - Raster scanning_ applying a dotmap image • c _ Observation of part buildingprocess through analysis of the II printing II programImproved part building through optimised process layoutFigure 1: Concept of investigating the effective mechanisms of MPCBesides process technological parameters of influence, the machine is of particular interest.

Essentially, the kinematic accuracy of the linear guides and the temperature dependent behaviour of the entire machine must be given in particular. Here, only the motion of the projector scanning the resin surface has direct effect on the exposure quality. Therefore, the influence of the accuracy of the scanner guides is of great importance. The scanner, driven by a step motor, moves with an178accuracy of 70Jlm per 90mm distance. In the direction perpendicular to the scan direction, the positioning accuracy is 20Jlm.The influence of the material on accuracy to form and size, also regarding geometric long term stability, is often underestimated. The two groups of photopolymers, acrylate- and epoxy based resins, have some very different characteristics (Jacobs, 1992, Xu, 1997/1). In this context, the influence of the material is not further followed up.

In this context, the influence of the material is not further followed up. These aspects are not a particular charac teristic of MPC but they are mentioned here because of their relevance for MPC.3 Experiments into Effective Mechanisms of the ProcessFor the experiments, all test samples were built using a test facility which has been developed for Multiple LED photographic curing at RIKEN. The material is an experimental acrylate-based photopolymer resin, HSX-V-2 (by Asahi Denka). At a constant vat temperature of 30°C, the resin has a viscosity of 11 =3 x I0-5Ns!cm2 • At the resin surface, the output power in each beam of light is PL,v =29.2JlW. The experiments aid the investigation of the characteristic features when curing a photopolymer using multiple beams of light.3.1 String FormationThe design of MPC necessitates simultaneous exposure of plane areas and thinnest stringswith the same constant scan velocity and LED power.

Therefore, in a computer-supported mathematical contemplation the exposure during curing using various numbers of beams of light for the exposure has been calculated. As a result, when exposing a plane area and a thinnest string at the same time, to ensure equal cure depth, the energy of at least five beams of light is necessary for exposing a single string. In this way, an equal level of exposure through superposition of sin gle exposures in one layer is always guaranteed. Only under these conditions, a thin string can be created while a plane area is cured at the same time without overcuring.For the initial experiments, single strings-exposed by five beam of light-have been chosenfor investigating the curing phenomena. A single string is of particular interest because of both their well..:known general importance when analysing the exposure process in Stereolithography in general, and investigating the characteristics of MPC in particular.

At scan velocities ranging from 0.1 to 1.6mm/s, each string has been exposed by one (LI) to five (L5) beams of light at a fixed hatch distance of 62.5Jlm. The beam diameter isO.5mm.Dimensions were measured using an optical microscope and a CCD (charge-coupled device) camera. A comparison of the line widths in scan direction and perpendicular to it at various scan velocities is shown in Figure 2. At con stant light power at the vat, PL,v =29.2JlW, it can be seen, that in general, the observed line widths decrease with increasing scan velocity. The graphs show, that in X-direction, the same line width can be achieved at higher scan speeds. Besides, perpendicular to scan direction line widths are slightly greater than they are in scan direction. The reason for that may be.a not correctly ad justed timer for turning on and off theLEDs. A change in the software should solve this issue. The influence of the different number of beams of light used for exposure is considerable.

A change in the software should solve this issue. The influence of the different number of beams of light used for exposure is considerable. At five beams of light, thinnest strings can be cured to up to 1.5mm/s and 1.3mm/s, respectively.179In X-direction1.2In V-direction1.0E..s J ..c....."0 .~E..s 0.8J ii 0.6 "0 .~III c;: :::::iIII c;: :::::i0.40.20.00.20.40.8 Scan velocity Vs [mm/s]0.61.01.21.41.60.00.20.40.8 Scan velocity Vs [mm/s]0.61.01.2Figure 2: Line width in X- and Y- direction at light power PL,v = 29.20IlWPerpendicular to the scan direction (Y-direction), at a scan speed of more than 1.3mm/s, nocuring occurs. Compared to the exposure in scan direction, to start the polymerisation perpen dicular to the scan direction, the scan velocity must be reduced, thus providing more exposure. The lower levels of curing found during exposure in Y-direction may be due to the exposure pat tern.

The lower levels of curing found during exposure in Y-direction may be due to the exposure pat tern. Across the scan direction, curing happens gradually and closely spaced. Initially, dots at a distance of O.5mm are exposed. The energy, however, is not sufficient to connect these individual dots on the resin surface. The next row of beams of light are, therefore, not able to expose the resin anywhere near the previously cured dots.

The next row of beams of light are, therefore, not able to expose the resin anywhere near the previously cured dots. As a result, no strings can be formed at all.Cured$trings - Lines are created bY drawing aStrings attached at one endsingle line with five beams of lightLine width increases with lower scanning velocitiesOn the free surface, curing dot-by dot requires more energy than at continuous exposure - Power PL v = 29.2r.tW - Material: acrylate resin HSX-V-2Vs = 1.5mm/sVs = 1.0mm/sVs = 0.5mm/sScan direction aVs = 1.3mm/sVs = 1.0mm/sVs = 0.5mm/sFigure 3: Formation of strings at various scan velocities vsin scan direction and across to it1801.41.6In scan (X-) direction, at scan velocities between 0.1 and 1.5mm/s, the shape of the stringsdiffers greatly (Figure 3). As expected, the line width increases with decreasing scan velocity. At 1.5mm/s, a string has a width of about 0.1 hnm.

As expected, the line width increases with decreasing scan velocity. At 1.5mm/s, a string has a width of about 0.1 hnm. Where the string is attached to a frame holding it in position in a right angle, the geometry is unclear. At low exposure, due to the low degree of polymerisation and its position free flowing on the resin surface, the string is rather soft and un stable. Xu et al. (1997/2) found that microscopic flow generated by the process heat could be observed. It is assumed that, during exposure, flow and probably shrinkage cause the string to move slightly on the resin surface. Width increases rapidly up to 0.68mm at 0.5mm/s. At 0.5mm/s, the string has been cured more and, therefore, is stronger and more stable. In Y-direction, at l.3mm/s, the string is cured at a width of about 0.31mm.

At 0.5mm/s, the string has been cured more and, therefore, is stronger and more stable. In Y-direction, at l.3mm/s, the string is cured at a width of about 0.31mm. At 0.5mm/s, its width increases to about 0.7mm.3.2 Connecting Mechanisms Between StringsThe connections between up to five adjacent lines formation a single string can be observed best at low scan velocities (Figure 4). The scan velocity is constant at O.lmm/s. When exposing with just one beam of light, the shape of the cured string reminds a little bit on a string of pearls. With increasing number of beams, the string becomes smoother.

With increasing number of beams, the string becomes smoother. At the same time, however, the line width increases only a little.Cured strings - Lines are created by drawing asingle line with one to five beams of lightShape of the line becomes smoother with increasing number of beams of light - Light power PL,V =29.21!W - Material: acrylate resin HSX-V-2 - When exposing one string with one beam of light, exposure happens in steps caused by the slightly jerky motion of the projector which is driven by a step motorExposure by orte beam of light05mm ~Exposure by two beams of lightExposure by three beams of lightWith increasing number of beams of light, this effect disappearsScan directionI:JFigure 4: Simultaneous multi-beam exposure at a scan velocity Vs =O.Imm1s in scan directionUneven curing when exposing at extreme low scan speeds is caused by the motion of theprojector.

In scan direction, the curing becomes sensitive to the slightest movements of the pro jector. The exposure happens in steps caused by the slightly jerky motion of the projector, which is driven by a step motor. This effect disappears with increasing number of beams of light expos-181ing the string. Each additional beam of light exposes the same string at a slightly shifted time ac cording to the arrangement of the beam array in the projector. Another observation is, at least three beams of light are necessary for curing one single string. Then, the contour of the string is similar to that of a straight line.3.3 Parameters of Influence "Energy Absorption"The course of the line widths and cure depths at various scan velocities and light power are shownillFigure 5. The data wereopta.ined after curing several strings with three beams of light. In X-girection, line width and cure .~epth increase with increasing light power. The same result can be found in Y-direction.

In X-girection, line width and cure .~epth increase with increasing light power. The same result can be found in Y-direction. However, the line width is in average slightly greater than in X direction. At scan velocities of 0.3 andO.4mmls, the increase in cure depth in Y-direction is slightl:y(steeper than that in X-direction. The reason for that may be the initially very low degree of curing at a low exposure dot-by-dot.

At higher exposure, the cure depths in X- and in Y direction are the same.Cured strings - Data obtained using an optical microscope and CCD cameraStrings were cured using 3-beam exposureIn average, line width in scan direction is slightly less than perpendicular to scan direction1.010.8J 0.6 ~ ,,;:0.4Q):.5 0.2Line width in X-direction1.010 ~ 0.4 :.5 0.28 .JO.6Q)Line width in V-directionIn average, cure depth in scandirection is slightly highter than perpendicular to scan direction - To get the highest scan velocity process can be operated at a power of 29.2 and 30.311W0.018.0 20.0 22.0 24.0 26.0 28.0 30.0 32.0 Light power Pev [f.lWlCure depth in X-direction1.00.018.0 20.0 22.0 24.0 26.0 28.0 30.0 32.0 Light power PLv!J,lW]Cure depth in V-direction1.0 !T'?T;;;rT'7TI'7;'R!Scan velocity:'-0.3mmls -X-0.4mmls -+-0.5mm/s -.-0.6mmls -0.7mm/s -·-0.8mm/s10.8u~ 0.6 .c 0.

~ 0.4 ~3 0.2 0.018.0 20.0 22.0 24.0 26.0 28.0 30.0 32.0 Light power Pev !J,lWl10.8 ,,]0.6 .c 0. ~ 0.4 230. 2Q)0.018.0 20.0 22.0 24.0 26.0 28.0 30.0 32.0 Light power PLV !J,lWjFigure 5: Parameters and their influence on curing3.4 Parameter of Influence "Part Geometry"The process parameters were varied over a wide range to find the optimum combination yielding good surface quality, fine detail, and sharp contours. The tests were carried out using the experimental resin HSX-V-2. The explanations in Figure 6 focus on the results achieved when building simple geometric objects. A mesh can be build with walls as thin as 0.2mm and in good accuracy in X-and Y-direction as well as at various angles. The effect of multi-spot exposure is182clearly visible in the vertical and the sloping elements. Part quality depends in the part orientation in the X-Y-plane. The vertical surfaces are smooth.

The effect of multi-spot exposure is182clearly visible in the vertical and the sloping elements. Part quality depends in the part orientation in the X-Y-plane. The vertical surfaces are smooth. It is one characteristic of MPC that strings which form layers lie exactly one above the other. This can not be changed since the arrangement of the lenses in the projector is fixed. MPC is successful for the process parameters for LED power and scan speed which were found above. A top view of the test part's surface shows that the lines from the exposure are clearly visible, even with the naked eye.Material - Acrylate resin H5X-V-2 - Resin viscosity 11 =3 x 10·sNs/cm2Mesh, top viewLayers, side viewProcess parameters - LED power PLY = 29.20/-lW - Wavelength 'A.

= 470nm - Scan velocity v, =O.75mm/sTest part - Size 45 x 45 x 9mm 3 - Build time t B = 81 minWall andsurlace, topv.iewCorner, top viewFigure 6: Influence of part geometry on surface texture and part quality4 DiscussionThe curing process in Multiple LED photographic curing (MPC) was experimentally exam ined. The effects of the parameters of influence on curing of single strings an formation of layers and three-dimensional objects were studied. Previously, theoretical investigations have shown that for the generation of equal cure depths in thin strings and plane areas, the exposure of at least five beams of light is necessary.The data reported in this paper show that strings can be accurately cured at low light power.Line widths increase at lower scan speeds. In scan direction, strings could be polymerised at a velocity of up to 1.5rnrn1s. At high scan velocities, the strings are unstable and the geometry is sometimes unclear.

In scan direction, strings could be polymerised at a velocity of up to 1.5rnrn1s. At high scan velocities, the strings are unstable and the geometry is sometimes unclear. One reason could be that microscopic flow causes the strings to move on the free resin surface. Besides, shrinkage may cause the strings to move. Across the scan direction, only up to 1.3rnrn1s strings could be created. These results can be explained by assuming that gradually and closely spaced exposure is not sufficient for polymerisation. At lower speeds, how-183ever, the exposed dots stick together and start formation strings. At extremely low scan velocity, the process becomes sensitive to slightest jerky movements of the projector when exposing with one beam of light. A jerky motion of the projector creates a line which looks like a string of pearls rather than a continuous line. Adding at least two beams leads to a smoother shape of the string.

A jerky motion of the projector creates a line which looks like a string of pearls rather than a continuous line. Adding at least two beams leads to a smoother shape of the string. At high velocities, this effect disappears completely.Due to the general characteristic of these investigations-strings were examined on a free resin surface-thisstudy·is limited to the basic curing process in MPC. It is the basis for further studies on the generation of three-dimensional objects, such as thin walls and other geometric shapes. The heatgeneration as the assumed reason for moves of the strings on the resin surface has notbeen examined and is subject of future examinations. The results of this study have impli cations/on further developments of the process and the design of the machine itself.ThIs study couldlead to new important findings about the MPC process.

The results of this study have impli cations/on further developments of the process and the design of the machine itself.ThIs study couldlead to new important findings about the MPC process. The results of this researchindicate thatMPCis a successful method for faithfullYigenerating three-dimensional ob jects,andhasthepotentialto aid design evaluation. Clearly, this technique has promise as a tool to fabricate three-dimensional objects in design evaluation. Future steps will thus include im provements in the light sources to increase process speed and accuracy of Multiple LED photo graphic curing.5 R.eferencesJacobs, P.F. (1992) RapidPrototyping and Manufacturing-Fundamentals of Stereolithography,Society of Manufacturing Engineers (SME), Dearborn, MILoose, K., Nakagawa, T., Niino, T. (1997) "Concept modelling using multiple LED photographic curing",proceedings, 13th Rapid Prototyping Symposium, Yokohama, Japan, pp. 35-40, No vember 1997.Marutani, Y., Kamitani, T.

FECTIVE THERMAL CONDUCTIVITIES OF UNIT-LATTICE STRUCTURES FOR MULTI-FUNCTIONAL COMPONENTSD. Cook*, S. Newbauer*, D. Pettis*, B. Knier† and S. Kumpaty, Ph.D., P.E.§Applied-Technology Center, Rapid-Prototyping Research, Milwaukee School of Engineering, Milwaukee, WI 53202 †MS Mechanical Engineering Student, UW Madison, Madison, WI 53202 §Department of Mechanical Engineering, MS Engineering Program Director, Milwaukee School of Engineering, Milwaukee, WI 53202AbstractApproaching the goal of automatically generating optimized multi-functionalcomponents, previously-identified unit-lattice structures are being characterized for their geometry-dependent, effective, thermal conductivities. This knowledge base will allow for the definition of low-mass, load-bearing, thermal-management structures. One application is a wearable power source for a custom, portable, active orthosis.

This knowledge base will allow for the definition of low-mass, load-bearing, thermal-management structures. One application is a wearable power source for a custom, portable, active orthosis. The function of this structure is to bear mechanical load while dissipating heat from the source, without burning the wearer. Additive manufacturing affords the capability of fabricating the resultant complex structures. Current research efforts are using finite-element analysis and physical testing to validate the characteristic models, and determining the scale dependence of internal-convective-flow development.

Internal-combustion engines, for example, average only 25% efficiency. For a 74.6kW (100Hp) engine, that is roughly 56kW of wasted energy, of which at least half is waste heat. That is 50MJ of wasted energy for a half-hour drive (maintaining average power)!The National Science Foundation’s (NSF) Engineering Research Center (ERC), titled theCenter for Compact and Efficient Fluid Power (CCEFP), for example, is working to improve system efficiencies of fluid-power systems at all levels, to increase the current average of 40%. This waste heat is not only an immediate loss of input power; it also reduces the viscosity of the oil, reducing the efficiency at which the system can use the power that it does receive. Cooling the oil from 90℃ to 80℃, using components with integrated passive thermal management, can improve system efficiency by 20% with minimal increase in mass [Herzog and Neveu, 2011].

Cooling the oil from 90℃ to 80℃, using components with integrated passive thermal management, can improve system efficiency by 20% with minimal increase in mass [Herzog and Neveu, 2011]. Generally, maintaining an operating temperature close to room temperature (at which components are designed by default) will minimize leakage and binding, thereby maximizing efficiency.Waste heat can also be converted to useful energy that can be fed back into the system, further improving system efficiency. For the above engine example, even at one percent (1%) conversion efficiency, thermoelectric generators could recover more than 280W to power electrical systems [Schneider et al., 2007]; and, efficiencies of 15-20% are anticipated using new materials [Canter, 2011].Finally, efforts to develop powered personal devices are hindered by the high temperatures generated, that can cause burning, pain or general discomfort.

Proper thermal management can mitigate these issues.The purpose of the research discussed herein is to contribute to a design paradigm, around the freedom afforded by additive manufacturing (AM), for multi-functional components and structures that: effectively manage waste heat, minimizing the detriment to the system (whilealso minimizing mass and bearing mechanical loads); improve safety; and, optimize the integration of devices for waste-heat recovery.B.Background: Considerable research has been conducted to characterize the thermal properties of cellular and lattice materials with the goal of establishing their benefits as thermal-management and multi-functional materials. Conventional heat sinks are not multi-functional, and generally limited in design by fabrication constraints.

Conventional heat sinks are not multi-functional, and generally limited in design by fabrication constraints. Predominantly, the most recent research involves foams, having generally isotropic properties, and homogeneous lattice structures, that do not conform to the application [Wadley, 2002].Metal foams are increasingly being used, since they can bear moderate loads in multipledirections, and are commercially available. This research takes this trend further to include anisotropic performances of devices for structural and thermal loads (as well as pressure-wave attenuation), and includes component integration, i.e. one component serving the functions of multiple conventional modular ones, making that one a true multi-functional component.Additionally, this prior research focused on forced-convective fluid flow through the material or structure ([Kim et al., 2004], [Seepersad et al., 2006], [Straatman et al., 2006]).

Contrary to the goal of developing multi-functional components that improve system efficiency, including mechanisms to provide this forced convection will detract from the efficiency gains.The scope of this research is in defining component geometries that provide greater heatefflux and, for personal applications, safe contact temperatures, while also contributing structurally. A portable, powered, ankle-foot orthosis is being developed within the CCEFP, designated as Test Bed 6 ([Human Assist Devices], [Shorter et al.]), for which the power source currently being developed is a small internal-combustion engine ([Project 2B.2], [Tian et al., 2010]). Certainly, there are concerns regarding the safety and noise of this device. The goal is to use the structure to protect the wearer from high temperatures, and provide optimal heat transfer for various components of the system.

Modeling in this way is also greatly simplified. This reference design may then also be used to initiate system optimizations that include such internal fluid flow.ApproachA.Unit-Lattice Characterization: Several unit-lattice structures had been previously identified for their applicability ingenerating structurally-optimized components ([Cook and Gervasi, 2011], [Cook et al., 2010]). Toward the goal of generating multi-functional components, three1 of these structures were characterized for their effective thermal conductivities as well. A unit cell is comprised of eight unit-lattice structures at the vertices of the hexahedral volume (Figure 1).Figure 1: (Left) unit lattice and (Right) unit cell of the Cube structure type.A thermal-resistance model had been developed to characterize the thermal properties ofcarbon foam [Yu et al. 2006]; and, a similar model was developed for the Cube unit cell [Appendix A].

2006]; and, a similar model was developed for the Cube unit cell [Appendix A]. Considering the complexity of the remaining structures, however, another means was required. The conductivity was, therefore, approximated using a rate equation known as Fourier’s law [Incropera & Dewitt, 1985]. Assuming that the temperature gradient across the unit cell is uni-directional (one-dimensional), the rate equation is expressed as:(cid:1869)"(cid:3051) (cid:3404) (cid:3398)(cid:1863)(cid:3032)(cid:3033)(cid:3033)(cid:3032)(cid:3030)(cid:3047)(cid:3036)(cid:3049)(cid:3032) (cid:1856)(cid:1846) (cid:1856)(cid:1876)⁄Where (cid:1869)"(cid:3051) is the heat flux in the x direction per unit area; keffective is the effective (bulk) thermal conductivity of the cell; and, (cid:1856)(cid:1846) (cid:1856)(cid:1876)⁄ is the temperature gradient across the cell, in the x direction.

Approximating the temperature gradient as ∆(cid:1846) ∆⁄ (cid:1876), and integrating both sides of the equation with respect to the cross-sectional area, A, Equation 3 becomes:(cid:1516) (cid:1869)"(cid:3051) (cid:1856)(cid:1827) (cid:3404) (cid:3398)(cid:1863)(cid:3032)(cid:3033)(cid:3033)(cid:3032)(cid:3030)(cid:3047)(cid:3036)(cid:3049)(cid:3032)(cid:1827) ∆(cid:1846) ∆(cid:1876)⁄The negative sign indicates that heat transfer is in the direction of decreasing temperature.COMSOL® provides the ability to calculate the heat flux using surface integration. Two opposing sides of the cube, in the direction of the desired thermal gradient, were set to a specific temperature differential; and, the other surfaces were considered insulated. The only unknown remaining is the effective thermal conductivity of the material. Initially, the fluid inside the structures is considered to be stagnant, so the effects of convection are neglected.

The only unknown remaining is the effective thermal conductivity of the material. Initially, the fluid inside the structures is considered to be stagnant, so the effects of convection are neglected. Effectively, the fluid is treated as an insulating solid, thereby creating a composite structure. Certainly, solid1 Only three of the five structures were analyzed due to time and budget constraints.(3)(4)insulators could be used as well, but few approach the low conductivity, mass and cost of air. The first analyses used the following parameters: Solid Material: Aluminum (k = 160 W/ m*K)  Fluid Material: Air (k varies with temperature, approximately = 0.025 W/m*K)  Cell Length (∆X): 0.0254 m (1 in)  Cell Surface Area (A): 6.45E-04 m2 (1 in2)  Temp Gradient (∆T) : 100KThe strut diameters, Ø, tested were 0.2L, 0.5L and 0.8L, where L is the edge length of theequilateral cubic cell.

With three different directions having three set diameters, 27 different combinations are possible; though, only 10 of them are “unique” combinations, assuming that orientation will not affect the outcome, i.e. Kxx of a Cube having Øx = 0.2L, Øy = 0.5L and Øz = 0.5L will be equal to Kyy of a Cube having Øx = 0.5L, Øy = 0.2L and Øz = 0.5L. The results of the conductivities for various cell sizing were then tabulated in Microsoft Excel® [Appendix C]; and, a regression tool (LINEST) was used to determine the equations relating the three orthogonal conductivities (Kxx, Kyy and Kzz) to the strut sizes (Øx, Øy and Øz) [Appendix E].Modifications to the Model:Initially, the fluid and solid on the sides of the unit cell were assigned the same temperature, since the distribution of heat flux was not known a priori; however, it was determined that this imposes an artificial temperature gradient and heat flux [Appendix B].

To remedy this, copper blocks were added fore and aft of the cell [Figure 2]. By placing the boundary conditions on the copper blocks, the heat flux is allowed to “choose its own path,” allowing for a more natural model.Figure 2: Inclusion of copper blocks in determination of effective thermal conductivity of the unit cell. As shown, the heat flux through the unit cell follows a natural path. For the 3-D structure, this results in non-planar temperatures at the faces of the cell, until the volume fraction of the solid approaches 100%.Because the conductivity of the copper blocks is known, the conductivity of the unit cell,at steady state, can be calculated using a one-dimensional thermal circuit, analogous to an electrical circuit, consisting of three resistors; two copper blocks, and the fluid-filled cell.

The “current” is the total heat flowing through the circuit, with the “voltage” being the temperature differential:∆(cid:1846) (cid:3404) (cid:1869)(cid:3051)(cid:1844)(cid:3047)(5)Where ∆T is the temperature difference from one node to another, qx is the total heatflowing through the circuit, and Rt is the overall thermal resistance. Because the model was set up as a series model, the effective resistance follows the relationship of:(cid:1844)(cid:3047) (cid:3404) (cid:1844)(cid:2869) (cid:3397) (cid:1844)(cid:2870) (cid:3397) (cid:1844)(cid:2871)Where R1, R2 and R3 are the respective resistors in the circuit. Solving for theconductivity of the fluid-filled unit cell is now just a matter of algebraic manipulation.

Solving for theconductivity of the fluid-filled unit cell is now just a matter of algebraic manipulation. Thermal resistance can be related to conductivity by:(cid:1844)(cid:3041) (cid:3404) ∆(cid:1850) (cid:1863)(cid:3041)⁄ (cid:1827)COMSOL® uses 400 W/m*K as the conductivity of copper, with the copper block being25.4 mm x 25.4 mm x 25.4 mm, resulting in a resistance of 0.0984252 K/W for each.B.Unit-Cell Convection: Initially, the air inside the structure was considered to be stagnant, so the effects ofconvection could be ignored; but, if the temperature difference across the cell becomes large enough, the fluid will start to move, driven by buoyancy forces due to a difference in density, generated by the temperature gradient. This is called “free” or “natural” convection.

This is called “free” or “natural” convection. In consideration of these cellular structures, two cases of enclosed natural convection will be discussed.If the temperature gradient occurs in the direction of the gravitational forces, two possiblesituations arise. One occurs when the density decreases in the direction of gravity. This will cause an unstable temperature gradient; and, the fluid will circulate. The second situation occurs when the density gradient acts in the opposite direction of gravity. This produces a stable temperature gradient; and, the fluid is stagnant.

Figure 4 shows the necessary characteristic length required for natural convection to occur, in relation to the temperature gradient across the cell.Cell Size vs. Temperature To Determine the Onset of Natural Convection25]m m[e z i S201510l l e C50300400500600700800900THot [K]deltaT = 100KDeltaT = 200kdeltaT = 300kFigure 4: Lower temperature differentials across the unit cells allows for larger cell sizes before the onset of internal natural convection.How will effective conductivity be affected by convection?Double-paned windows are used in homes because the air film acts as an insulator; however, if convective currents develop inside a window, the total heat transfer across the window will increase. It was initially expected that similar results would occur with the cubes. To test this, models were set up in COMSOL®.

It was initially expected that similar results would occur with the cubes. To test this, models were set up in COMSOL®. The setup was similar to that of earlier tests, with a unit cell structure and adjacent copper blocks, and temperatures set at the ends of the copper. A local volume force, Fb, was applied to the fluid using the density calculated by the(5)1000non-isothermal-flow solver to approximate the buoyancy force created when the fluid heats up and expands:(cid:2162)(cid:4652)(cid:4652)(cid:3377)(cid:3029) (cid:3404) (cid:3398)(cid:2025)(cid:3041)(cid:3036)(cid:3047)(cid:3033) (cid:1499) (cid:2189)(cid:4652)(cid:4652)(cid:3377)Where ρ is the density of the fluid at a given node, and g is the gravitational constant (9.81 m/s2). Again, only cubic volumes were analyzed; limiting the number of analyses to be conducted, and matching the expected cubic nature of the resultant multi-functional structure.C.

Again, only cubic volumes were analyzed; limiting the number of analyses to be conducted, and matching the expected cubic nature of the resultant multi-functional structure.C. Initial finite-element analysis (FEA) StepsConductivity Optimization:In order to design a thermal-management structure (TMS) with a desired temperaturedrop, the scale of the thermal conductivity must be known. A 2-D axi-symmetric structure was designed within COMSOL®. The cylinder had a hollow inner radius, representing the heat source. The inner and outer radii of the structure were selected; and, a hemi-spherical top was added to enclose the heat source. The model was then divided into subsections, each assigned material properties that closely resembled those of insulators. A temperature was prescribed at the inner radius of the structure; and, natural convection at room temperature was specified at the outer radius of the component. The model was solved for the temperature at the outer surface.

The model was solved for the temperature at the outer surface. The subsections were then redefined with new material properties until a desired temperature drop was achieved. This was a manual iterative optimization, but, will be automated in future work. Once the desired temperature drop was achieved, the scales of the thermal conductivities for each subsection were noted. These were then used in designing the preliminary structure.Preliminary Structure DesignThe bulk thermal conductivity of the unit lattice structures had been determined as well as the target thermal conductivities of the structure design. To design the preliminary structure with the derived unit-cell, geometry-dependent, effective-conductivity relations, the subsections of the structure had to be kept as cubic as possible.

To design the preliminary structure with the derived unit-cell, geometry-dependent, effective-conductivity relations, the subsections of the structure had to be kept as cubic as possible. Geometric relations determining the distance of each strut from the center of the cylinder and height of each subsection were derived based on the; number of radial components, number of columns, and radii of the structure. A script was written to calculate these values, varying the radial components and columns. The thermal conductivity could then be calculated for each section, varying the diameter of the strut.The smallest diameter achievable with our current casting procedure is one millimeter(1mm). It was determined that, even with the smallest diameter strut, the target thermal conductivities could not be achieved. The initial bulk-thermal-conductivity equation derived was based on aluminum as the material of the Cube.

The initial bulk-thermal-conductivity equation derived was based on aluminum as the material of the Cube. New equations were derived by changing the material of the Cube in the FEA. Comparing the derived equations showed that the original equation can be scaled to fit the new equations by multiplying the original with the factor (cid:1863)(cid:3041)(cid:3032)(cid:3050) (cid:3040)(cid:3028)(cid:3047)(cid:3032)(cid:3045)(cid:3036)(cid:3028)(cid:3039) (cid:1863)(cid:3042)(cid:3045)(cid:3036)(cid:3034)(cid:3036)(cid:3041)(cid:3028)(cid:3039) comparing the calculated bulk thermal conductivity of the original equation and the FEA results of the new material. The target thermal conductivity could then be achieved by using Mold Star 22™, a material with a thermal conductivity of 43.2 W/mK, to cast the one-millimeter-diameter struts.⁄.

The target thermal conductivity could then be achieved by using Mold Star 22™, a material with a thermal conductivity of 43.2 W/mK, to cast the one-millimeter-diameter struts.⁄. This resulted in a maximum of seven percent (7%) error when(6)The preliminary structure was chosen to have 5 columns and 15 radial components (Figures 6,7). The hemi-spherical top part of the structure was not designed based on the derived bulk-thermal-conductivity equation, but was approximated to match the desired conductivities while maintaining good structural connectivity to the lower cylindrical section. A 3-D structure was generated and cast, as was a finned thermal-management device of the same material and total mass for reference (Figure 7). The engineered structure had twice the surface area of the finned component.FabricationThe one-millimeter (1 mm) minimum feature size on a structure of this size and complexity is quite a challenge to produce.

The weight of the alloys is also of concern. Certainly, direct-metal processes are an option, but were not used due to limitations of time and funding, as well as uncertainty of the thermal properties of the additively-manufactured metal.The selected method for fabricating the thermal-management structure was to use additive manufacturing to produce a pattern for investment casting. This approach provides the potential for investigating the performance of many alloys and composites. Indeed, while the investigations to date have included exclusively air-filled structures, the principles will be extended to include solid fill. This is of particular interest for combining wear resistance and heat dissipation in the structural design of components.The patterns for the thermal-management devices were created in nylon at the Milwaukee School of Engineering (MSOE), using Selective Laser Sintering™ (SLS™), and then encased in ceramic.

After burning out the nylon, the empty ceramic shell was filled with the molten Moldstar 22™ alloy, using MSOE’s proprietary high-resolution investment-casting process.ResultsA.Unit-Lattice Characterization: The resultant unit-cell thermal resistances and conductivities are tabulated in AppendixC. Multiple relation forms were attempted in the regressions, including: quadratic; second-order cross-multiplied; third order; second-order quadratic; linear cross-multiplied; linear; and, homogeneous second-order [Appendix D].

Comparing the errors of these relations in predicting the performance of the unit cells, it was found that the second-order cross-multiplied form had the lowest error; however, this is not definitive: LINEST is limited to 17 terms; and, efforts were constrained by time.Effects of ConvectionInitial analyses that included the development of internal convective flow returnedunexpected results: heat transfer in the orthogonal directions showed negligible change. The majority of the heat was flowing through the solid. This makes sense if thought about as an electrical model. Each Cube can be thought of as three resistances in parallel, solid, fluid and a series combination of solid and fluid [Yu et. al. 2006]. Even if the total resistance of the fluid resistor decreased, e.g. by changing the fluid, the thermal resistivity is still substantially larger than that of the solid metal used.

al. 2006]. Even if the total resistance of the fluid resistor decreased, e.g. by changing the fluid, the thermal resistivity is still substantially larger than that of the solid metal used. This raised the question of, “What would happen if the conductivity of the solid decreased, thereby increasing its thermal resistance?”To show the relationship between material and convection, a new model was used,consisting of a block of air, surrounded by a block of metal. By varying the thickness of the metal surrounding the air, a plot of percent difference in heat transfer (between convective and stagnant air) versus solid thickness could be made for various materials. The plot generated shows that as the conductivity of the material increases, or as the amount of solid material around the air increases, the effects of convection becomes less prominent (Figure 5).

The plot generated shows that as the conductivity of the material increases, or as the amount of solid material around the air increases, the effects of convection becomes less prominent (Figure 5). But, even with a material of low conductivity, such as Inconel, the percent difference between convective and stagnant air is low (0.6% at a 100K delta). The exponential increase in the significance of the internal convection, with decreasing solid volume fraction and conductivity, shows that the solid material must have thermal conductivity close to that of the contained fluid for cellular convective flow to be of significance, relative to the stagnant conductivity.Increase in Heat Flux Through Convection vs.

Solid Volume Fraction]0.7%[x u l F t a e H ni0.60.50.40.3e c n e r e f f i D %0.20.100510152025Volume Fraction of Solid [%]Inconel (k=11.2 W/m_K) Aluminum (k=200 W/m_K)Moldstar (k=43 W/m_K)Figure 5: This plot shows the dependence of the heat flow rate through the unit cell on the development of cellular convective flow, at a ΔT = 100K. Even for solid materials of low thermal conductivity, and at low volume fraction, the change in heat flow is less than one percent in comparison to stagnant air. Convective flow through a lattice structure, however, will produce a significant difference in the heat flow rate.B.Structure: A Matlab® script was composed to assist in the design of the preliminary thermal-management structure, incorporating the derived geometric relations to maintain “square” unit cells, calculating the resultant outer radius of the structure as a function of the number of angular and radial divisions (Figure 6).

For the one-millimeter-minimum-diameter constraint, a large outer radius of 40 mm was selected, corresponding to fifteen angular and five radial divisions. Certainly, with higher-resolution fabrication processes, more compact (smaller radius) designs can be achieved.30Outer Radius vs. Slices and DivisionsFigure 6: Outer radius of the thermal-management structure (vertical axis [mm]), constrained by “squareness” of the resulting cells. The X-axis represents the number of angular divisions, while the Y-axis represents the number of radial divisions. Choosing 15 slices (angular divisions) and 5 layers (radial divisions), the cell sizes in each layer were determined from the underlying relation.For the Moldstar 22™ material, this provided effective conductivities across the radialdivisions of keffective = {6.4371, 2.5359, 0.9432, 0.3113, 0.0731}.

With all necessary parameters defined, the model was generated, fabricated and cast, along with a “reference” finned heat sink (Figure 7).Figure 7: (Left) engineered thermal-management structure, (right) finned heat sink of equal mass and material, but half of the surface area.C.Experimental Results The engineered structure, finned component, and a cylindrical piece of aluminum foamwere tested for their thermal performance in providing a safe contact temperature. A heating element was placed inside each structure. The voltage to the heating element was controlled with a variable supply; and, both voltage and current were measured. A type-K thermocouple was used to record the temperature of the structure. Measurements were taken after five minutes of supplying the measured power to the heating element, for six consecutive power settings.

Measurements were taken after five minutes of supplying the measured power to the heating element, for six consecutive power settings. The measured temperatures were then plotted as functions of the input power (Figure 8).The engineered structure performed exceptionally better than the finned component andaluminum foam. The finned component outperformed the aluminum foam, but both were harmful to touch after the second measurement.e r u t a r e p m e T e c a f r u S e r u t c u r t)C g e D(200 180 160 140 120 100 80 60 40 20 0Experiment: Comparison of Thermal-Management StructuresEngineered StructureAluminum FoamFinned Heat Sink"Pain Threshold""Toasted Skin"S05 20 Power Into Heat Source (W)101525Figure 8: Experimental results of testing the engineered structure, aluminum foam and a finned heat sink.

Only the engineered structure provided a safe contact temperature.It should be noted that the foam structure was not of the same material, nor of the same mass as the engineered structure or finned component: an off-the-shelf cylindrical foam piece was used; but, its size was close to that of the finned structure. The small pore sizes of the aluminum foam, in comparison to the cavities in the engineered structure, also did not allow for any significant fluid flow through the component, or radiation. Fluid flow through the engineered structure, driven by buoyancy, and radiation through the radially-expanding cavities, allowed the structure to maintain a safe contact temperature even past its original design specification.D.Varying the Design: 2-D axi-symmetric FEA models were analyzed, varying the number of columns in theanalysis from three to seven. The results show that, with fewer columns, the structure is able toachieve a greater temperature drop.

The results show that, with fewer columns, the structure is able toachieve a greater temperature drop. Even though a structural analysis was not performed on the current design, fabricating a structure with less supports could lead to a fragile component.The model was also evaluated with Inconel (thermal conductivity approximately 11.2W/mK). The results showed that, with the current geometry, Inconel could achieve a temperature drop over three times that achieved with Mold Star 22™. This means that using Inconel as the material of the component could lead to more compact, and stronger, components. The rate of heat dissipation will drop, however, increasing the core temperature for a given heat production rate. This balance requires optimization in the design.Strut diameters smaller than one millimeter were also investigated. For Inconel structures, an engineered TMS with a strut diameter of 0.5 mm had twice the performance of a structure with a strut diameter of 1 mm.

For Inconel structures, an engineered TMS with a strut diameter of 0.5 mm had twice the performance of a structure with a strut diameter of 1 mm. Again, Inconel had nearly three times the performance of Mold Star 22™ at strut diameters of 0.5 mm.Improvements to the current design can easily be made by varying the material of thecomponent. The components can also have increased effectiveness and greater compactness if manufacturing techniques are improved upon.Next StepsOf greatest priority is the application of this research to an integrative thermal-management structure for the powered orthosis. The engine is expected to be operational shortly; and, testing on patients is anticipated in the summer of 2012.

The engine is expected to be operational shortly; and, testing on patients is anticipated in the summer of 2012. In addition to protecting the wearer from the excessive temperatures, investigations are already underway to determine the practicality of integrating thermoelectric generators into the structure to power some of the electrical components, thereby improving system efficiency.For this work, it was assumed that the lower surface temperatures and higher heat-fluxrates of the engineered thermal-management structure, due to convective fluid flows through it, further improves the performance of the component; therefore, no effort was made to account for convective through-flow. To achieve precise metrics, and full understanding of the structure’s performance, future research will investigate the fundamental dynamics that both drive and resist this through-flow, including: Nusselt number – ratio of convective/conductive heat transfer.

 Darcy flow relation to define resistance to convective through-flow.  Schmidt number – the ratio of momentum diffusivity (viscosity) and mass diffusivity, used to characterize fluid flows in which there are simultaneous momentum and mass- diffusion convection processes.To expedite the analyses, a slice of the structure is being used, with “symmetry”boundary conditions on the cut faces (Figure 9).Figure 9: This is a vertical slice of a revolutionary-generated structure that is being used for analyses of the performance of the resultant structure. “Symmetry” boundary conditions are applied to the cut faces to reflect the missing geometry.

“Symmetry” boundary conditions are applied to the cut faces to reflect the missing geometry. A similar structure will be fit around the small internal-combustion engine on the powered orthosis to dissipate heat and protect the wearer from prolonged exposure to high temperatures.Ten geometric configurations of each unit-lattice cell were analyzed, providing a total of27 data points for each through the assumption of symmetry and negligible influence of fluid convection. As with the structural characterizations done in the past, more configurations could be analyzed for greater resolution or fidelity, particularly in the case of very small solid geometries.The second-order, cross-multiplied form was selected as the best representation, out ofthe seven that were considered, of the geometry-dependence of the unit-lattice’s effective thermal conductivity, but is certainly not the only applicable form.

Simpler relations may meet the tolerance of the application, and may even be more accurate.While work has begun to determine the effective thermal conductivity of these unit- lattice structures in directions not aligned with their orthogonal members, further research is required to determine the extent of their effective-conductivity anisotropy. For those cases where internal natural convection can be neglected, the thermal characteristics are expected to be orthotropic, just as their structural properties are, with constant factors that relate to the orthogonal directions, e.g. the fraction of the x-direction conductivity that contributes to the diagonal.Future efforts will also investigate the potential for automating the optimization of the distribution of effective conductivity throughout the thermal-management structure.

This will allow for higher resolution and better performance, while allowing the designer to consider more options as well.ConclusionsPassive thermal management of components can be achieved through the design of components using engineered lattices. To this end, the geometry-dependent effective bulk thermal conductivities of three unit lattice structures were determined for the three orthogonal directions. The heat flux through the unit cell scaled linearly to its size, resulting in a constant conductivity, simplifying the design of gradient lattice structures.While internal cellular convection may develop and alter the effective bulk conductivity, the investigations showed that the effects were negligible for the materials under consideration.

Generally, the target performance of the thermal-management structure can be attained by varying the base material and geometry.The engineered thermal-management structure outperformed aluminum foam and anequal-mass finned component in providing a safe contact temperature. Convective flow of air through the final structure, however, did alter the performance significantly. While this improved the safety of the device, it does present a challenge in accurately predicting performance. This is being addressed in current research efforts.Benefits of this approach to thermal management include the potential for creating integrated, minimal-mass, multifunctional components, and for optimizing temperature gradients for energy recovery through thermoelectric generation.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEFFECTIVENESS OF EX-SITU HEAT TREATMENT OF L-PBF AM 17-4PH STAINLESS STEEL SPECIMENS INTENTIONALLY EXPOSED TO DIFFERENT AS-BUILT THERMAL HISTORIESC. R. Hasbrouck 1,*, Joseph Bartolai 2, Darren C. Pagan 3, and Simon W. Miller 21 Department of Industrial and Manufacturing Engineering 2 Applied Research Laboratory 3 Department of Materials Science and Engineering The Pennsylvania State University, University Park, PA 16802AbstractThe presented research demonstrates the effectiveness of H900 heat treatment in eliminating microstructural and mechanical property differences between additively manufactured 17-4PH stainless steel samples of varying thermal histories.

For this effort, 17-4PH stainless steel was manufactured using laser-based powder bed fusion on an EOS M280 machine in two geometries: ASTM Standard E8 subsize rectangular tensile specimen geometry with thicknesses of 2mm and 6mm. Thermal histories were manipulated by adding secondary laser passes on each layer at varying levels of reduced power. All samples were heat treated after the build following H900 procedures. Mechanical performance was evaluated with uniaxial quasi-static tensile testing and Vickers microhardness measurements. Metallography was examined qualitatively with optical and electron microscopy as well as quantitatively through electron backscatter diffraction.

Metallography was examined qualitatively with optical and electron microscopy as well as quantitatively through electron backscatter diffraction. No statistically significant mechanical property or microstructural differences were discovered, suggesting a successful ex-situ heat treatment.Keywords: laser powder bed fusion, 17-4PH stainless steel, thermal history, heat treatmentIntroduction and MotivationSince its development in the 1980s, additive manufacturing (AM) has been an exploding area of research due to its ability to build parts in a layer-by-layer fashion based on slices taken from three-dimensional data [1]. This layer-by-layer buildup allows for greater geometric freedom and minimizes material waste when compared to traditional subtractive manufacturing methods such as machining. Additionally, since AM requires no tooling, it allows for expedited, low-cost prototyping when a design needs multiple iterations in a short time frame.

Additionally, since AM requires no tooling, it allows for expedited, low-cost prototyping when a design needs multiple iterations in a short time frame. AM has more recently gained traction as a popular method for producing metallic parts with complex geometries for critical applications within the aerospace and medical industries [2]. However, all current metallic AM systems use either powder metals, solidification processing, or a combination of both that results in metallurgical differences between AM components and conventional wrought or cast products [2].

Research involving the measurement and mitigation of such AM attributes as mechanical anisotropy, residual stress, and defects is critical for maintaining and predicting the performance of metallic parts produced by AM [3].Corresponding author: crh33@psu.edu388Of the seven overarching categories of AM, typically only three are used for directly producing metallic components: powder bed fusion (PBF), binder jetting, and directed energy deposition (DED) [1]. Binder jetting is less popular than PBF or DED for critical components due to the need for a post-process infiltration or sintering process. These post-processing steps may lead to localized chemistry and mechanical performance differences, or porosity and dimensional instability, respectively. Both PBF and DED build layers on a solid metal substrate and frequently use either a laser or an electron beam as the energy source for melting the metallic feedstock, typically powder.

Both PBF and DED build layers on a solid metal substrate and frequently use either a laser or an electron beam as the energy source for melting the metallic feedstock, typically powder. Electron beams are typically less popular energy sources because they require the build volume to be under vacuum and require magnetic lenses that limit the builds to non- magnetic metals [4]. Both PBF and powder-based DED use powder as the feedstock, but PBF uses a powder bed and DED uses a powder feed [5]. Powder-based DED is popular for additive repair of worn or damaged components, but does not have the level of dimensional control that PBF does [5]. Additionally, DED can use a wire feedstock and a plasma arc power source for faster builds, but sacrifices additional detail and resolution in the process [2].

Additionally, DED can use a wire feedstock and a plasma arc power source for faster builds, but sacrifices additional detail and resolution in the process [2]. Therefore, laser-based powder bed fusion (L-PBF) is a popular AM method for producing a wide range of metallic alloys with complex geometries, excellent resolution, and relatively easy-to-manipulate build parameters.Stainless steels are popular for AM processes due to their combination of good corrosion resistance and achievable mechanical properties. In particular, 17-4PH (AISI 630) stainless steel is a valuable alloy for applications up to 600°F where corrosion resistance and high strength are necessary [6], such as aerospace, nuclear, naval, and chemical industries [7]. AM 17-4PH steel has been compared to its wrought counterparts for development of ex-situ heat treatment practices [6], [8], [9], cyclic deformation and fatigue behavior [7], and stress corrosion cracking [10].

In addition to more typical applications of AM 17-4PH, for example a complex geometry hydraulic manifold [11], AM 17-4PH has also been explored for its high temperature and/or high strain rate performance [12], [13]. AM 17-4PH has been used to produce thin walled parts, where part placement [14] and laser scan length [15] were explored for their effects on resulting microstructure. Additionally, part location [16], build orientation [17], and scan pattern [18] have been explored for their effects on AM 17-4PH fatigue behavior and other mechanical properties, while laser defocusing has been examined for effects on resulting defects and microstructure of AM 17-4PH [19]. In addition to variations in microstructure and mechanical response of 17-4PH produced by AM, its corrosion resistance may be effected by such AM parameters as build bed location [20].

In addition to variations in microstructure and mechanical response of 17-4PH produced by AM, its corrosion resistance may be effected by such AM parameters as build bed location [20]. Like many other AM metallic alloys, 17-4PH has been studied with and without hot isostatic pressing post-build [21], and in both as-built and heat treated conditions.In addition to the intentional manipulation of build parameters for L-PBF, it is important to note that the chemical composition of the feedstock powder can have a significant effect on the as-built microstructure and properties, as well as the post-build heat treatment response of AM 17- 4PH [8], [22], [23]. This is most evident for nitrogen-atomized 17-4PH powder that retains enough nitrogen, a strong austenite stabilizer, to greatly increase the amount of retained austenite (RA) during solidification [24].

The alloy 17-4PH is considered a martensitic precipitation hardening stainless steel [25], so any RA remaining in the as-built condition is expected to transform into martensite during heat treatment. If the austenite is stabilized by a high presence of nitrogen, a fully martensitic matrix may not be achieved upon traditional heat treatment regimens [8].389The typical reference for heat treating this alloy in its wrought form is ASTM Standard A564, “Standard Specification for Hot-Rolled and Cold-Finished Age-Hardening Stainless Steel Bars and Shapes,” [26]. The highest strength achievable for this alloy occurs at the peak aging condition, designated as H900. Note that the numbers following the initial H in these precipitation hardening thermal cycles indicates the temperature at which the alloy is aged. For example, H900 indicates an aging temperature of 900°F, H1100 indicates an aging temperature of 1100°F, and so on.

For example, H900 indicates an aging temperature of 900°F, H1100 indicates an aging temperature of 1100°F, and so on. The heat treatment associated with H900 has two main steps: a solution anneal and an aging cycle. The hold time for each step is typically one hour per inch of material thickness. The purpose of solution annealing is to form a fully austenitic microstructure, which involves raising and holding the temperature of the samples above the austenite start temperature of the alloy, or ~1300°F for 17-4PH [27]. This austenization cycle effectively “erases” any prior thermal history from manufacturing by dissolving the constituent elements [25]. The solution anneal is followed immediately by either an oil (sections up to 3in) or forced air (sections ≥ 3in) quench to room temperature, forming a metastable low-strength martensitic microstructure supersaturated with copper [25].

The second step of H900 involves aging the alloy to peak strength by heating and holding the samples at the aging temperature of 900°F, forming copper precipitates that increase ultimate tensile strength to a peak of ~190ksi [25]. The H900 heat treatment has been used in other AM 17-4PH studies [9], [17], [22], [23]. Outside of the H900 peak aging heat treatment, the H1100 [7], [12] and H1150 [22] overaging heat treatments are popular for the AM 17-4PH alloy due to the associated increases in corrosion resistance and fatigue performance over the H900 highest-strength condition. Additional references for heat treatment of 17-4PH that may be used include the ASM Handbook Volume 4D, Heat Treating of Irons and Steels [25] and the ASM Metals Handbook Desk Edition [28].

Additional references for heat treatment of 17-4PH that may be used include the ASM Handbook Volume 4D, Heat Treating of Irons and Steels [25] and the ASM Metals Handbook Desk Edition [28]. Furthermore, some studies have been performed that explore new heat treatment methods for AM 17-4PH outside of conventional wrought standards [6], [8].low-carbonWhile the microstructure of AM 17-4PH has been examined in the literature both as-built and after a number of different thermal cycles, the alloy (whether conventionally or additively manufactured) is nearly always tested for mechanical response after heat treatment to take advantage of the precipitation hardening behavior. Relatively common mechanical property tests used for AM 17-4PH include quasi-static tension testing [8], [12], [16]–[18], compression testing [12], [13], [21], and cyclical fatigue testing [7], [16], [17].

In addition to AM 17-4PH characterization, quasi-static tensile testing is a highly utilized destructive mechanical property test for both fundamental scientific research studies as well as manufacturing process qualifications for most metallic alloy systems in all types of applications, and therefore should be well understood by those who work in manufacturing of metals. The ubiquitous nature of this test is likely not only due to its simplicity, but also because it generates material property data for more than one mechanical response. A single uniaxial tension test can directly measure such material property data as yield strength, ultimate tensile strength, and elongation/ductility [29]. It can also provide indirect calculated estimates of resilience and toughness based on the directly measured properties. An illustration of a typical stress-strain curve from a uniaxial tension test for a ductile material is shown in Figure 1.

An illustration of a typical stress-strain curve from a uniaxial tension test for a ductile material is shown in Figure 1. In addition to the marked strength values, the resilience of the sample is estimated by the area under the stress-strain curve corresponding to the linear elastic region and the toughness is estimated by the area under the entire stress-strain curve.390Ultimate Tensile StrengthσFracture StrengthFracture X, s s e r tYield StrengthS g n i r e e n i g n EElastic Modulus (slope of linear region)Elastic deformationUniform plastic deformationNon-uniform plastic deformation (necking)Engineering Strain, εFigure 1: Generic stress-strain curve from a uniaxial tension test of a ductile material.Because mechanical properties are directly impacted by microstructure, that is also an important consideration when evaluating AM 17-4PH.

Popular characterization methods for microstructure include x-ray diffraction (XRD) for phase identification [6], [9], [13], [18], [19], [23], especially when RA is suspected, and electron backscatter diffraction (EBSD) for grain size, morphology, and orientation [7], [9], [10], [14], [15], [17], [21].

Additional characterization methods for AM 17-4PH seen in the literature include microhardness mapping [18] and fracture surface examination [7], [8], [10], [16]–[18] for mechanical response, and optical [6]–[8], [12], [13], [17]–[19], [21], [23], scanning [6], [8]–[10], [15], [19], [21], and transmission electron microscopy [9], [15], [19] for microstructural and chemical composition analysis.Although mechanical response and microstructural development have been studied extensively in the literature for L-PBF AM 17-4PH samples, the authors believe this is the first research that intentionally modified thermal history of 17-4PH parts by adding additional laser exposures to the material after initial melt and solidification of the feedstock powder. The effectiveness of a standard H900 heat treatment was then verified by mechanical property testing and microstructural characterization.

The effectiveness of a standard H900 heat treatment was then verified by mechanical property testing and microstructural characterization. As in the literature, quasi-static tension testing and EBSD were used to analyze mechanical response and microstructure, respectively, after heat treatment. Additionally, microhardness mapping was used to examine local properties, and qualitative metallography was performed by both optical and electron microscopy techniques.

The remainder of the paper details the research questions answered by this study; the experimental methodology including: powder distribution and chemistry analysis, machine details and build parameters, heat391treatment procedures, mechanical testing setups, and metallographic procedures; the quantitative data analyses, and the conclusions drawn from this work.Research QuestionsTwo research questions (RQ) and an associated hypothesis were developed to understand the effectiveness of ex-situ heat treatment at eliminating microstructural and mechanical response differences of laser-based powder bed fusion (L-PBF) additively manufactured (AM) 17-4PH stainless steel specimens intentionally exposed to different as-built thermal histories:RQ1: Can a H900 heat treatment effectively standardize the microstructural and mechanical responses of a L-PBF AM 17-4PH stainless steel specimen exposed to a traditional L-PBF AM as-built thermal history?RQ2: Can a H900 heat treatment effectively standardize the microstructural and mechanical responses of a L-PBF AM 17-4PH stainless steel specimen exposed to additional reduced power laser passes during build?Hypothesis: Based on previous research in both AM of 17-4PH stainless steel and more traditional wrought or cast processing, the authors predicted that a H900 heat treatment will effectively standardize the microstructural and mechanical responses of L-PBF AM 17-4PH stainless steel specimens intentionally exposed to these different as-built thermal histories.Experimental ProceduresThe experimental procedures will be expanded upon in subsections for each major step of the research, including building the specimens, heat treatment procedures, mechanical testing procedures, and microstructural characterization.Building the SpecimensThe feedstock powder used for this research was AMP 17-4Ar (argon atomized) stainless steel powder supplied by Powder Alloy Corporation (PAC).

The composition of this powder follows the chemical specifications of UNS S17400, ASTM A705, and ISO 15156-3 alloy compositions [30]. The specific chemical composition details as measured by NSL Analytical, an ISO17025 approved test laboratory, are detailed in Table 1.392Table 1: Chemical composition of 17-4PH stainless steel feedstock powder.Element Fe Cr Ni Cu Nb+Ta C Mn Si S P O NSpec. Min. Results Spec. Max. Balance 15.88 4.24 3.19 0.25 0.03 0.43 0.74 0.01 0.02 0.1 < 0.1Balance 15.00 3.00 3.00 0.15Balance 17.00 5.00 5.00 0.45 0.07 1.00 1.00 0.03 0.04 0.1 0.1Particle size distribution and flow characteristics were measured by both PAC and the Center for Innovative Materials Processing through Direct Digital Deposition (CIMP-3D) at Penn State. PAC measured volumetric particle size distribution using a sieve test according to ASTM Standard B214 [31].

PAC measured volumetric particle size distribution using a sieve test according to ASTM Standard B214 [31]. CIMP-3D measured both numeric and volumetric particle size distributions using a TESCAN MIRA high-resolution analytical scanning electron microscope (SEM) utilizing light scattering per ASTM Standard B822 [32]. Both PAC and CIMP-3D measured the Hall flow rate of the powder samples per ASTM Standard B213 [33]. The results of these measurements are detailed in Table 2.Table 2: Particle analysis results for size distribution and Hall flow rate.D10 (µm) D50 (µm) D90 (µm) Hall Flow Average (s/50 g)PAC (Volumetric) CIMP-3D (Volumetric) CIMP-3D (Numeric) PAC CIMP-3D26.76 25.41 1.5638.06 3.6454.76 50.33 34.9715.7 12.05The large difference in the numeric and volumetric particle size distributions reported is due to the presence of a large number of fine particles.

The volume based particle size distribution as measured by CIMP-3D is illustrated in Figure 2.393Figure 2: PAC 17-4PH stainless steel volume based particle size distribution.The presence of fine particles is further illustrated by the SEM images shown in Figure 3. It can be also seen from these images that some of the powder particles experienced collisions with other molten particles during argon atomization and that some particles are extremely lacking in sphericity.(A)(B)Figure 3: 17-4PH feedstock powder particles as imaged with SEM. (A) 750X magnification, and (B) 1500X magnification. Arrows indicate examples of powder particles that collided during atomization.394All parts used for this research were built on an EOS M280 laser-based powder bed fusion (L-PBF) machine at CIMP-3D. The parts studied included two geometries of interest, namely ASTM Standard E8 [29] subsize rectangular tensile specimen geometry with thicknesses of 2mm (“thin”) and 6mm (“standard”).

The parts studied included two geometries of interest, namely ASTM Standard E8 [29] subsize rectangular tensile specimen geometry with thicknesses of 2mm (“thin”) and 6mm (“standard”). Additional cylindrical parts of 8mm diameter and the same height (100mm) were built and exposed to the same thermal history conditions as the tensile specimens. These cylinders were cut into disks of ~3.75mm thickness post-build for dynamic testing using a Split Hopkinson pressure bar (SHPB) setup.Two separate builds were created with identical build layouts and sample numbering schemes, aside from the initial build number identifiers. Both builds used standard EOS parameters for 17-4PH stainless steel, including: laser power of 220W, hatch spacing of 0.11mm, scan speed of 755.5mm/s, beam offset of 2µm, and layer thickness of 40µm.

Both builds used standard EOS parameters for 17-4PH stainless steel, including: laser power of 220W, hatch spacing of 0.11mm, scan speed of 755.5mm/s, beam offset of 2µm, and layer thickness of 40µm. In addition to standard building parameters with single full power laser exposures at each build layer, two unique build parameter combinations were explored with the tensile and SHPB sample geometries: double exposure and mixed exposure. Double exposure indicates that the full Z-length of the sample was exposed to secondary laser passes at reduced power for each build layer. Mixed exposure indicates that the bottom half of the Z-length (0-50mm) of each sample was exposed to single full power laser passes, and the top half of the Z-length (50-100mm) of the sample was exposed to additional secondary laser passes at reduced power for each layer. This was performed to create a sharp gradient between as-built thermal histories.

This was performed to create a sharp gradient between as-built thermal histories. A schematic illustrating the differences between standard, double, and mixed exposure is shown in Figure 4.Z- (Build) directionPass 1: Full Power Pass 2: Reduced PowerSingle Double MixedFigure 4: Build parameter combinations for single, double, and mixed exposures.For build 1, the secondary laser passes were performed at 50% (“half power,” 110W) of the standard parameters’ laser power. For build 2, the secondary laser passes were performed at 25% (“quarter power,” 55W) of the standard parameters’ laser power. Each build included three replicates of each tensile thickness (standard and thin) for each build laser exposure combination (single, double, and mixed), and two replicates of the SHPB cylinders at each build laser exposure combination. It should be noted that these parts were built in conjunction with cylindrical lattice geometries used for a different research project.

It should be noted that these parts were built in conjunction with cylindrical lattice geometries used for a different research project. Figure 5 highlights an example build layout. As the key indicates, samples marked in gray were built with single full power laser exposures at each layer, parts marked in green had double laser exposure (half power for build 1, quarter power for build 2), and parts marked in purple experienced mixed laser exposure. A completed build is shown in Figure 6.395SingleDoubleMixedSHPB 43-2 3-1SHPB 62-2 2-1SHPB 5Thin 3-3 Thin 3-2 Thin 3-1SHPB 3Thin 1-3 Thin 1-2 Thin 1-1SHPB 25-2 5-1Y1-2 1-1SHPB 14-2 4-1Thin 2-3 Thin 2-2 Thin 2-1XFigure 5: Example build plate layout and naming scheme.

Circles indicate cylindrical lattice parts that were used for a separate research study.396Figure 6: A completed build on the substrate.Heat Treatment ProceduresThe parts used in this study were subjected to the peak aged, highest-strength H900 heat treatment per ASTM Standard A564 [26]. The specifics of the H900 treatment for these builds are: • Parts all heat treated while remaining attached to the substrate to prevent geometricdistortion per ASTM Standard F3301 [34]Parts heat treated in a standard air furnace atmosphere per ASM Handbook Volume 4D: Heat Treating of Irons and Steels [25]Solution annealo Temperature: 1900°F ± 25°F o Hold time: 1 hour +15 / -0 minutes o Nitrogen cooled to below 90°FAgeo Temperature: 900°F ± 10°F o Hold time: 1 hour +15 / -0 minutes o Nitrogen cooled to below 90°F397Mechanical Testing ProceduresMechanical testing was performed to characterize both global and local mechanical response.

For global properties, quasi-static uniaxial tensile testing was performed. For local properties, microhardness mapping was used. Additionally, qualitative examination of global mechanical response was completed via fracture surface analysis. Each procedure is expanded upon separately.Tensile Testing ProceduresAll tensile specimens were built and tested according to ASTM Standard E8 [29] using subsize rectangular tensile specimen geometry (flat dogbones) with thicknesses of either 2mm (“thin”) or 6mm (standard). The samples were tested on Instron 8801MT servo-hydraulic frame equipped with an Instron 22kip load cell. The tests were conducted in displacement control with a constant X-head rate of 0.05in/min, which translates to a nominal strain rate within the linear elastic region of 0.013in/in/min. An externally mounted Instron 2620-824 axial extensometer with a 40% full-scale strain range was also used.

An externally mounted Instron 2620-824 axial extensometer with a 40% full-scale strain range was also used. The testing setup is shown in Figure 7.Figure 7: Quasi-static tensile testing specimen and machine setup.398During all specimen testing, data for tensile force via the load cell, and axial strain via the extensometer was collected at 10Hz. The tensile strain was calculated using input dimensions for the cross-sectional length and width of the gauge region as measured with dial calipers pre-testing. The data was then plotted using Microsoft Excel to determine ultimate tensile strength, elastic modulus, yield strength (via 0.2% linear offset method), and strain at fracture.

The data was then plotted using Microsoft Excel to determine ultimate tensile strength, elastic modulus, yield strength (via 0.2% linear offset method), and strain at fracture. Statistical analysis of the data across thermal history conditions was performed using Microsoft Excel; more information on the results and statistical analysis are in the Results, Analysis, and Discussion section.Microhardness Mapping ProceduresThe microhardness mapping was performed on specimens intended for dynamic compressive testing using the Split Hopkinson pressure bar (SHPB) setup.

Due to the similar cross- sectional areas between the tensile specimens (36mm2 in the gauge length) and the SHPB cylinders (50mm2), plus the identical heights between them (100mm) it can be assumed that the thermal histories are similar enough for direct comparison.Four replicates of each thermal history condition (standard, double exposure at half power, and double exposure at quarter power) were each hot mounted together with glass-filled epoxy powder. This resulted in three plastic pucks, each containing four individual samples, for a total of 12 microstructural characterization samples. The SHPB disks were mounted across the XZ/YZ plane such that the surface being examined included a cross-sectional view across the build direction layers. The samples were mechanically polished to a surface finish of 1µm with an automated polishing machine using standard metallographic procedures per ASTM Standard E3 [35].

The samples were mechanically polished to a surface finish of 1µm with an automated polishing machine using standard metallographic procedures per ASTM Standard E3 [35]. A schematic of the sample mount used for this testing is shown in Figure 8.Figure 8: Representative sample mount with four replicates of a given thermal history. The diamonds represent the microhardness indentation map pattern, the gold dotted lines indicate where along the build height the exposure strategy was changed for mixed specimens, and the white arrows indicate the build (Z-) direction.399All microhardness measurements were performed according to ASTM Standard E92 for Vickers hardness [36]. Using a Wilson VH3100 Vickers automated hardness tester, ~550-600 indents in a rectangular grid were performed per sample with a 500g load, 10s dwell time, and 0.1778mm (0.007in) spacing. The testing setup is shown in Figure 9.

The testing setup is shown in Figure 9. The resulting microhardness maps are in the Results, Analysis, and Discussion section.Figure 9: Microhardness mapping setup.Fracture Surface AnalysisThe fracture surfaces of the tensile specimens were examined with a TESCAN MIRA high- resolution analytical SEM using a secondary electron detector. A 10.0kV accelerating voltage under high vacuum was used, and micrographs were taken at various magnifications until 5000X (a view field of 55.4µm) was established as an appropriate representation of the fracture behavior. Representative fracture surface micrographs are documented in the Results, Analysis, and Discussion section.Microstructural CharacterizationMicrostructural characterization was performed both qualitatively and quantitatively.

All microstructural analysis was performed on the specimens intended for dynamic compressive testing using the Split Hopkinson pressure bar (SHPB) setup as described in the Microhardness Mapping Procedures section. For qualitative analysis, both optical and electron microscopy was400performed on etched samples. For quantitative analysis, electron backscatter diffraction (EBSD) measurements were used. Each procedure is expanded upon separately.Qualitative MetallographyThe same samples and sample preparation techniques as described in the Microhardness Mapping Procedures section were used for qualitative metallographic analysis. Upon completion of polishing to a surface finish of 1µm, the samples were immersion etched with Kalling’s Reagent for 15-20s. Representative micrographs at 800X were taken both optically with a Keyence VHX- 6000 microscope and via secondary electron imaging with a TESCAN MIRA high-resolution analytical SEM at 10.0kV accelerating voltage.

Representative micrographs are documented in the Results, Analysis, and Discussion section.Quantitative MetallographyThe same samples and sample preparation techniques as described in the Microhardness Mapping Procedures section were used for qualitative metallographic analysis. Upon completion of polishing to a surface finish of 1µm, an overnight vibratory final polish to a surface finish of 0.02µm was performed prior to electron backscatter diffraction (EBSD) measurements.EBSD measurements were performed on a ThermoScientific Apreo S field emission scanning electron microscope (FESEM).

The following settings were used for EBSD measurements:Specimen tilt: 70° • Magnification: 2000X (≈ 135µm by 200µm scanning area) • Beam current: 3.2nA • Accelerating voltage: 20.0kV • Step size: 0.5µm • Possible phases: body-centered cubic (bcc) iron and face-centered cubic (fcc) iron • Grain boundary threshold: 10° misalignmentPost processing of data was completed with Oxford Instruments’ AZtec software, Microsoft Excel, and MATLAB. Data of interest from AZtec included grain area minimum, maximum, mean, and standard deviation, as well as ASTM grain size and grain equivalent circle diameter. Excel was used to add 95% confidence intervals on the mean grain size and grain equivalent circle diameter. MATLAB was used for data visualization and the full quantitative statistical analysis using logarithmic grain size distributions.

MATLAB was used for data visualization and the full quantitative statistical analysis using logarithmic grain size distributions. More information on the results and statistical analysis are in the Results, Analysis, and Discussion section.401Results, Analysis, and DiscussionThe related results for each type of mechanical testing or microstructural analysis will beexpanded upon as subsections in this section for clarity.Tensile Testing ResultsFour global mechanical properties were analyzed to determine the effectiveness of heat treatment normalizing different L-BPF thermal histories: 0.2% offset yield strength (YS), elastic modulus, ultimate tensile strength (UTS), and strain at break. Figure 10 shows a summary of each of these tensile properties measured through quasi-static uniaxial tensile testing. The color of each data point indicates which laser scan strategy was used to produce that specimen.

The color of each data point indicates which laser scan strategy was used to produce that specimen. The shape of the data point indicates from which build each specimen was produced. Visually, there appears to beFigure 10: Scatter plots of tensile testing results. (A) shows 0.2% yield strength, (B) shows elastic modulus, (C) shows UTS, (D) shows strain at break. Changes in build parameters are indicted with different colors. Shape of the data point indicates build number.402no significant difference in mechanical response between the different scan strategies, which would indicate that H900 is an effective heat treatment for L-PBF produced 17-4PH parts.To confirm this observation, statistical analysis for the mechanical testing data was completed using Microsoft Excel.

Thermal history and thickness combination categories included: standard 6mm, double half 6mm, double quarter 6mm, mixed half 6mm, mixed quarter 6mm, standard 2mm, double half 2mm, double quarter 2mm, mixed half 2mm, and mixed quarter 6mm. Mechanical testing metrics of interest included the 0.2% offset yield strength, elastic modulus, ultimate tensile strength, and elongation at break. The sample size in most categories prohibited a normality test, so normality was assumed for this analysis. Next, an F-test was completed for each category compared to standard to verify equivalence of variances. Finally, a paired two-tail t-test was performed compared to standard to assess the equivalence of the means of each sample category. The results for each of these tests is seen in Table 3 for the 6mm thick samples and Table 4 for the 2mm thick samples.

T-tests comparing each scan strategy with standard parameters for either the 6mm or 2mm thick specimen report p-values above 0.05, indicating a 95% confidence that post heat treatment, all of the specimen come from populations with identical means. There is one exception, elastic modulus of the 2mm mixed half power specimens. Because the other mechanical properties for this same set of specimens do not report a statistically significant difference, and the small sample size, this one result is considered to be an outlier.It is also important to note the fracture location for each of the tensile specimens to identify any potential localized differences in mechanical response. For example, the largest difference in thermal history of any samples should occur in the mixed exposure specimens at the build’s halfway transition in Z-length (50mm) where the exposure settings changed from standard single exposure to a double exposure with either half or quarter laser power.

It follows that if the as-built thermal history remained after heat treatment, the samples should have all broken near the middle of the gauge length where that transition occurred. Figure 11 and Figure 12 below show that none of the mixed exposure samples, with either half power (Figure 11) or quarter power (Figure 12) secondary exposures, failed at the center of the gauge length for either specimen thickness.Some kind of text here to fix my figure numbering(A)(B)Figure 11: Fractured tensile specimens with mixed exposure, where the bottom half of the sample received standard single laser exposure and the top half of the sample received half laser power secondary exposure. (A) 6mm standard width specimens, and (B) 2mm thin width specimens.(A)(B)Figure 12: Fractured tensile specimens with mixed exposure, where the bottom half of the sample received standard single laser exposure and the top half of the sample received quarter laser power secondary exposure.

(A) 6mm standard width specimens, and (B) 2mm thin width specimens.404However, upon further inspection, there does appear to be a pattern in the location of where the mixed exposure samples fractured. When the samples are laid out such that all of the halves that were attached to the substrate are on the same side, every mixed exposure specimen fractured in the region of the gauge length corresponding to the standard parameter settings. This is shown in Figure 13. It should be noted that for tensile testing, the half of the tensile specimen that was attached to the substrate was randomized for whether it was placed in the stationary or moving machine grip.(A)(B)Figure 13: Mixed exposure samples aligned by halves closest to the substrate. (A) Mixed half power specimens, and (B) mixed quarter power specimens.For comparison, all double and standard exposure samples fractured in a more randomized fashion, and no obvious patterns or correlations to build direction were found.

This is illustrated by the samples shown in Figure 14 and Figure 15 for double exposure and standard exposure, respectively. It is possible that the change in exposure patterns halfway through the part changed the local mechanical properties, but the global properties remained the same due to breaking in the weaker half of the sample. Additionally, the properties of the specimen with multiple laser exposures throughout the specimen did not exhibit different mechanical properties than the specimens with a mixed exposure scan pattern. More investigation of local properties and microstructural differences in the tensile specimens is needed to confirm if there is a local change in mechanical properties at the scan pattern change boundary. If this is the case, it is unclear why the double exposure samples did not show a change in global mechanical properties due to the difference in exposure across the entire length of the sample compared to the standard exposure samples.

More experimentation would be required to verify this.405(A)(B)Figure 14: Double exposure samples aligned by halves closest to the substrate. (A) Double half power specimens, and (B) double quarter power specimens.(A)(B)Figure 15: Standard exposure samples aligned by halves closest to the substrate. (A) Standard 6mm specimens, and (B) standard 2mm specimens.Microhardness Mapping ResultsLocal mechanical response was explored using the SHPB samples. Four replicates of each scan strategy were examined with microhardness mapping.

Four replicates of each scan strategy were examined with microhardness mapping. The resulting maps are shown in Figure 16 for single exposure (A), alternating half exposure (B), and alternating quarter exposure (C) scan strategies.406(A)Replicate 1Replicate 2Replicate 45mmReplicate 3(B)Replicate 1Replicate 4Replicate 2Replicate 35mm(C)Replicate 1Replicate 4Replicate 2Replicate 35mmFigure 16: Microhardness mapping results for (A) single exposure, (B) alternating half exposure, and (C) alternating quarter exposure.407The color scale in Vickers hardness (HV) was selected as the full range of Rockwell C hardness (HRC) values to highlight the homogeneity across the samples, rather than modifying the scale to show larger differences in local microhardness that may sway the interpretation of the results to indicate that there were significant differences in local properties across the scan strategies.

Each of the replicates from the scan strategies shows a local microhardness equivalent to ~41.7-49.7 HRC, which is consistent with the expected average hardness after peak aging of ~45 HRC. Some inhomogeneity is expected with L-PBF due to pores and other defects inherent to the process.Fracture Surface Examination ResultsA 10.0kV accelerating voltage under high vacuum was used to examine all of the tensile specimens’ fracture surfaces Micrographs were taken using secondary electrons at 5000X (a view field of 55.4µm). Each of the scan strategies in both thin and standard geometries exhibited fracture surfaces typical of ductile failure.

Representative tensile testing fracture surfaces for standard full power exposure (A), double exposure at half power (B), and double exposure at quarter power (C) are shown in Figure 17.(A)(B)(C)Figure 17: Representative tensile testing fracture surfaces for specimens built with (A) standard full power exposure, (B) double exposure at half power, and (C) double exposure at quarter power.Qualitative Microstructural Characterization ResultsRepresentative optical (A) and SEM (B) micrographs of standard single full powerexposure samples etched with Kalling’s Reagent at 800X are shown in Figure 18.408Figure 18: Representative optical (A) and SEM (B) micrographs of standard single full power exposure samples etched with Kalling’s Reagent at 800X.The micrographs illustrate the presence of a fully martensitic matrix with the presence of large, dispersed carbides throughout.

Energy dispersive x-ray spectroscopy (EDS) mapping was not performed to assess the allying element forming the carbides, but it is anticipated to be niobium. Additionally, imaging at a high enough magnification to see the copper carbides was not possible with the optical technique and was therefore also skipped for the SEM. The EDS mapping for carbides and copper precipitates should be explored in future work.Qualitative Microstructural Characterization ResultsWhile collecting EBSD data, several maps are available to visualize the data. For this particular study, the Z-direction inverse pole figures (IPF) coloring overlaid with grain boundaries provide the best visual representation of the microstructural features that were being measured, including grain size, grain morphology, phase identification, and grain orientation.

All samples measured (four replicates of each thermal history condition) resulted in phase identifications of 100% body-centered cubic (bcc) iron, indicating that no retained austenite was present and the sample microstructural matrices were completely aged martensite. The grain size measurements from AZtec reported by area are shown in Table 5. Similarly, the grain morphology measurements from AZtec reported by equivalent circle diameter are listed in Table 6. Note that the grains that intersected the border of the image area were not counted for either measurement.

Note that the grains that intersected the border of the image area were not counted for either measurement. Both tables include an “average” row that considers the results of all four replicates in each thermal history category.409Table 5: Grain measurement by area in µm2 as reported by AZtec.Grain Measurement by AreaGrain Count 1049 1073 1154 1090 "Average" 1091.5Category Replicate1 2 3 4StandardMixed HalfMixed Quarter907 1028 972 947 "Average" 963.5 989 936 1224 9431 2 3 41 2 3 4Mean Area (µm2) 20.3 20.6 17.5 19.6 19.5 22.4 19.6 20.9 22.0 21.2 19.7 22.4 16.8 20.3 19.8Min Area (µm2) 2.5 2.5 2.5 2.5 2.5 2.5 2.5 2.5 2.5 2.5 2.5 2.5 2.5 2.5 2.5Max Area (µm2) 483.3 751.3 380.5 554.8 542.5 425.5 506.5 557.5 694.8 546.1 337.3 474.0 329.5 729.3 467.5Std. Dev.

Dev. (µm2) 37.4 41.8 29.2 42.4 37.7 42.1 39.9 40.9 49.1 43.0 33.8 42.5 29.4 44.5 37.6"Average" 1023.0Table 6: Grain measurement by equivalent circle diameter in µm as reported by AZtec.Grain Measurement by Equivalent Circle Diameter Std. Dev.

The build direction is indicated by the white arrow; note that since all four replicates for each thermal history condition were hot mounted together in a single 1.5” diameter specimen, it was not always possible to have +Z in the same direction and still have the required distance from the sample to the detector. Each micrograph has approximate dimensions of 135µm by 200µm. The black pixels correspond to no identified solutions for the proposed phases (bcc iron and fcc iron) in any orientation. The solution rate was 90% or higher for all measured samples.411(1)(2)ZZ(3)(4)ZZFigure 19: Z-direction inverse pole figure colored maps overlaid with grain boundaries for standard exposure specimens. The numbers in the upper right-hand corners correspond to the replicate number.412(1)ZZ(3)ZZFigure 20: Z-direction inverse pole figure colored maps overlaid with grain boundaries for mixed half exposure specimens.

The numbers in the upper right-hand corners correspond to the replicate number.413(2)(4)(1)ZZ(3)ZZFigure 21: Z-direction inverse pole figure colored maps overlaid with grain boundaries for mixed quarter exposure specimens. The numbers in the upper right-hand corners correspond to the replicate number.414(2)(4)Microstructural Characterization Analysis and DiscussionThe statistical analysis for the microstructural characterization data was completed using MATLAB. Thermal history categories included: standard exposure (“single power”), mixed alternating standard and half power exposure (“half power”), and mixed alternating standard and quarter power exposure (“quarter power”). The microstructural features of interest included phase identification, grain size by area, and grain morphology by equivalent circle diameter.

The microstructural features of interest included phase identification, grain size by area, and grain morphology by equivalent circle diameter. Since all phases measured were bcc, no statistical analysis was necessary to confirm that the phases were equivalent across all samples, so only grain area and equivalent circle diameters were analyzed further. Due to the nature of grain size measurements, a logarithmic distribution was used instead of a normal distribution for this data analysis.First, to better visualize the data distribution, normalized histograms for each of the three categories, including all four replicates, were created for both the grain areas and grain equivalent circle diameters.

These can be seen in Figure 22 and Figure 23, respectively.Figure 22: Log grain area normalized histograms by thermal history category.415Figure 23: Log grain equivalent circle diameter normalized histograms by thermal history category.Next, a Kolmogorov–Smirnov (K-S) test was performed to determine whether all of the replicates for each thermal history category could be combined into one corresponding dataset per category for both grain area and equivalent circle diameter. The K-S test serves as a nonparametric evaluation of the equality of continuous, one-dimensional probability distributions that can be used to compare two or more samples.

The K-S test serves as a nonparametric evaluation of the equality of continuous, one-dimensional probability distributions that can be used to compare two or more samples. The K-S tests (total of 36) for each thermal history category for both grain area and equivalent circle diameter determined that all replicates can be combined and treated as one overarching parameter set.For additional data distribution visualization and an opportunity to determine significant difference in means by inspection, normalized box and whisker plots were created for both the grain area (Figure 24) and grain equivalent circle diameter (Figure 25) data across each thermal history category with combined replicate data.416Group 1: Single Group 2: Half Group 3: QuarterFigure 24: Log grain area normalized box & whisker plot by thermal history category with combined replicate data.

Left (Group 1) is the normalized box and whisker plot for the grain area of the standard thermal history builds, middle (Group 2) is for the mixed alternating standard and half power exposure samples, and right (Group 3) is for the mixed alternating standard and quarter power exposure samples.Some text here to fix my figure numbering.Group 1: Single Group 2: Half Group 3: QuarterFigure 25: Log grain equivalent circle diameter normalized box & whisker plot by thermal history category with combined replicate data. Left (Group 1) is the normalized box and whisker plot for the grain equivalent circle diameter of the standard thermal history builds, middle (Group 2) is for the mixed alternating standard and half power exposure samples, and right (Group 3) is for the mixed alternating standard and quarter power exposure samples.417Finally, analysis of variance (ANOVA) was used to test for differences between two or more means by thermal history category.

All six ANOVAs (three for grain area and three for grain equivalent circle diameter) resulted in no statistically significant difference between means across thermal history categories. Therefore, no quantitative difference in resulting microstructure across thermal history categories was discovered with EBSD analysis.ConclusionsThis research suggests that an ex-situ H900 heat treatment is effective in eliminating microstructural and mechanical property differences between 17-4PH stainless steel samples produced with drastically varied L-PBF AM thermal histories. Standard quasi-static tensile testing results revealed no statistically significant differences in global mechanical response through yield strength, ultimate strength, or elongation at break across the five thermal history categories for each thickness subset (2mm and 6mm) of specimens. Additionally, no significant differences in local mechanical response as examined by microhardness mapping were observed.

Additionally, no significant differences in local mechanical response as examined by microhardness mapping were observed. Qualitative metallography indicated ductile fracture across all tensile specimens, and no discernable differences in microstructural features such as matrix phase or precipitation features were observed with optical or SEM techniques. Similarly, EBSD measurements revealed no statistically significant differences in microstructure between samples as identified by grain areas, grain equivalent circle diameters, or metallurgical phase fractions. The results of this research are encouraging when considering the possibility of extrapolation to homogenizing any differences in microstructure and mechanical properties of other metallic alloys produced by AM processes with inherent thermal history differences by implementation of ex-situ heat treatments.

However, because no analysis was performed on as-built specimens prior to heat treating, it is currently impossible to conclude whether the results are due to an effective heat treatment or ineffective processing parameters.To further investigate potential differences in local versus global mechanical response and corresponding microstructure, several characterization techniques are proposed for future work. Using digital image correlation (DIC) during tensile testing could help examine localized strain behaviors. For local microstructural differences, using an EBSD map that stitches several scan areas would help to identify phases and grain sizes and morphologies, and electron probe microanalysis (EPMA) could be used to detect any potential differences in chemical composition due to loss of certain alloying elements during solidification.

Energy dispersive x-ray spectroscopy can be used to map the compositions of the various precipitates that form at different scales during the solidification and heat treatment of this alloy. Additionally, it would be valuable to use a design of experiments to examine the effects of scan patterns, build plate layout, and tool paths on the localized mechanical response and microstructure. Finally, all of the work already completed here and proposed for future study of these H900 parts should be replicated on the as-built versions of these samples to see what the pre-heat treatment mechanical response and microstructure would be both locally and globally to get a full understanding of how intentionally manipulating thermal histories affects L-PBF AM 17-4PH parts.418DisclaimerThis material is based on work funded by the Defense Threat Reduction Agency (DTRA) through the Naval Sea Systems Command (NAVSEA) under Contract No. N000024-18-F-6401, Delivery Order 21F8353.

N000024-18-F-6401, Delivery Order 21F8353. The opinions, findings, conclusions, and recommendations expressed in this material are those of the authors and do not necessarily reflect the views of NAVSEA, nor do they reflect the views of DTRA or collaborators at NAVSEA. Additional funding for this work was provided by the Materials Research Institute (MRI) at Penn State’s Seed Grant 2020 program.AcknowledgmentsThe authors would like to thank Dr. Ken Meinert and Mr. Corey Dickman from the Center for Innovative Materials Processing through Direct Digital Deposition (CIMP-3D) at Penn State for performing the powder distribution analysis and assistance with building the metal samples used in this study, respectively, as well as Mr. Eric Strauch of the Applied Research Laboratory (ARL) at Penn State for performing the tensile testing of the specimens.

Eric Strauch of the Applied Research Laboratory (ARL) at Penn State for performing the tensile testing of the specimens. An additional acknowledgment is due to CIMP-3D for allowing access to their metallographic laboratory for sample preparation and the Materials Characterization Laboratory (MCL) at Penn State for equipment time on the vibratory polisher and FESEM for EBSD measurements. Finally, the support from Solar Atmospheres in Hermitage, PA with heat treating the samples and Acura-Cut in Pleasant Gap, PA for wire EDM services is greatly appreciated.References[1] I. Gibson, D. Rosen, and B. Stucker, Additive Manufacturing Technologies, 2nd ed. NewYork, NY: Springer, 2015.[2] J. O. Milewski, Additive Manufacturing of Metals. Springer International Publishing, 2017.doi: 10.1007/978-3-319-58205-4.[3] T. DebRoy et al., “Additive manufacturing of metallic components – Process, structure andproperties,” Prog. Mater. Sci., vol. 92, pp. 112–224, Mar.

viewed PaperEFFECTS DUE TO VARIATIONS IN THERMAL PROPERTIES OF MARAGING STEEL FOR FINITE ELEMENT MODELING OF THE WAAM PROCESSMatthew Register*† and Matthew W. Priddy*†Department of Mechanical Engineering, Mississippi State University, 210 Carpenter Hall, Mississippi State, MS 39762, USA †Center for Advanced Vehicular Systems, 200 Research Blvd, Starkville, MS 39759, USAAbstractFinite element (FE) thermal simulations of the wire arc additive manufacturing (WAAM) process have been widely used to predict the temperature history of as-built parts. Temperature-dependent thermal properties like density, conductivity, specific heat, and latent heat are required to accurately simulate the solidus to liquidus transition seen in the heat affected zone.

Current research has shown that thermal properties measured experimentally or simulated using material database software can be used for thermal modeling; however, there has been no direct comparison shown to determine which is most appropriate for WAAM modeling. The focus of this research is to compare the temperature variation of the FE thermal simulations with experimentally measured and computer-generated properties for M250 grade maraging steel. The thermal history for thin wall builds are compared with differing temperature-dependent thermal properties to examine the relationship between thermal properties and history. It was shown that constant values at room temperature result in an increased thermal response for WAAM thermal simulations and linearized properties decreased the temperature when both were compared with the experimental thermal property inputs.

Further work needs to be performed before determining the most appropriate source for thermal properties (wrought, wire, as-built, or material database), but it can be concluded that the form of material does affect the thermal response for maraging grade 250 steel.Introduction and BackgroundWire arc additive manufacturing (WAAM) is a metal based additive manufacturing (AM) technology used to produce large near net shape parts [1], [2]. The process utilizes in the form of an arc to melt wire feedstock in a layer-by-layer fashion [3], [4]. Standard welding equipment is often used such as gas metal arc welding, gas tungsten arc welding, or plasma arc welding. WAAM has been an increasing area of research in AM due to its high deposition rates and material efficiency compared to other metal AM processes like laser powder bed fusion [1], [3], [5].

WAAM has been an increasing area of research in AM due to its high deposition rates and material efficiency compared to other metal AM processes like laser powder bed fusion [1], [3], [5]. With deposition rates up to 10 kg/hr [5], it can be used to manufacture large components in a reasonable time compared to traditional casting methods. However, parts manufactured using WAAM often experience challenges caused by the large heat input and thermal cycling introduced into the part during printing.Repeated heating and cooling cycles during the printing process leads to non-uniform thermal expansion and contraction of the material inducing residual stresses and part distortion [1], [6]. The WAAM process is intended for large-scale parts; however, with increased part size there is a greater chance of geometric inaccuracies due to a reduction in cooling rate as the part height increases [7].

As a result, the accuracy of the thermal history of a printed part is essential in determining the resultant microstructure, mechanical properties, and geometrical accuracy [8]. FE analysis can be used to predict the temperature history and residual stress of the WAAM process before building the part experimentally, saving time and labor costs [9]–[13].The FE thermal simulation requires the use of temperature dependent thermal conductivity, specific heat, and latent heat of fusion; however, there has been little evidence supporting which source is best for these properties regarding WAAM simulations. Additionally, temperature dependent thermal properties can be very difficult to find in literature depending on the amount of research on the material of interest. In literature, properties which cannot be found are often generated by input of the elemental composition into a material database software [14], [15].

In literature, properties which cannot be found are often generated by input of the elemental composition into a material database software [14], [15]. Constant property values associated with room temperature (RT) have also been used as oversimplifications when the properties cannot be found in literature [16]. Experimentally measured1436values can be collected using digital scanning calorimetry (DSC) and laser flash technique (LFA);however, this can be costly and time consuming [17], [18]. To add to this variety, the properties can be gathered across a large range of temperatures from RT to well above the melting temperature. Lastly, it is unclear how the material state could play a role in thermal response.

Lastly, it is unclear how the material state could play a role in thermal response. For example, the filler wire used for the welding process may result in thermal properties that differ from wrought or as-built material.This paper investigates the variability in temperature history associated with variety of the thermal property inputs in a FE WAAM simulation using Abaqus/Standard 2019. Temperature dependent properties investigated include thermal conductivity, specific heat, and latent heat of fusion for filler wire and as-built material. Maraging grade 250 (M250) was chosen as the material of interest due to its excellent weldability and mechanical strength [19]. Several cases were tested including the actual measured values determined by DSC and LFA, constant values at RT, and linearized properties.

Several cases were tested including the actual measured values determined by DSC and LFA, constant values at RT, and linearized properties. The nodal temperature was simulated for a WAAM thin wall build, and values were measured at the center of the part on the top surfaces of layers 0, 1, 5, and 10 to best understand how various thermal property inputs can affect the temperature history in the part.Materials and MethodologyExperimental Build DetailsA 10-layered thin wall was built with a 1.1 mm diameter M250 wire using a Fronius TPS 4000 Cold Metal Transfer (CMT) Advanced Welding Package. The material used for the substrate was A36 mild steel. Machine data was recorded throughout the deposition process to be used later for the simulation to accurately match the experimental conditions. The gathered data includes voltage, current, travel speed, x, y, z coordinate locations, and arc activation. The printing process parameters used are listed in Table 1.Table 1.

The gathered data includes voltage, current, travel speed, x, y, z coordinate locations, and arc activation. The printing process parameters used are listed in Table 1.Table 1. The WAAM process parameters used during the build. Current [A] Voltage [V] Power [W] Travel Speed [mm/s] Wire Feed Speed [in/min]16015240010250With this information, the print path was modeled precisely to match the experimental setup. During the deposition process, four thermocouples (type K, size 0.032” Omega model KMQXL-032U-24) were used to measure the temperature variation over time. The locations of each thermocouple can be seen in Figure 1.Figure 1. The CAD drawing for the thin wall (top left), the built part (bottom left), the thermocouple locations (right) are shown for the RTRC thin wall print.The thin wall was measured to be 80 x 6.5 x 20.6 mm (L x W x H) and the substrate is 101.6 x 101.6 x 12.5 mm (L x W x H).

The location of the thermocouples is important during FE thermal model calibration to effectively capture the temperature gradient through the substrate. The experimental data will be used to compare against the FE thermal simulation results once the proper thermal property input is determined. Although the FE simulation is overpredicting the temperature past layer five, it has good agreement in the first five layers. It was decided this was appropriate based on similar data from calibrated results from other FE modeling papers [11],1437[18], [20]. The sharp peaks in the experimental data are believed to be caused either from shielding gas effects or electrical impedance from the CMT arc, but the true cause has not been determined.

The sharp peaks in the experimental data are believed to be caused either from shielding gas effects or electrical impedance from the CMT arc, but the true cause has not been determined. In future modeling efforts, the cooling rates will need to be adjusted after the first five layers most likely to have better agreement to the experimental data.FE Thermal ModelThe FE thermal simulation was built using Abaqus/Standard 2019 and the AM modeler plugin. The heattransfer found in WAAM is driven by the transient heat conduction Equation 1 described below:(1)𝑑𝑑𝑑𝑑 𝐶𝐶𝑝𝑝𝜌𝜌 𝑑𝑑𝑑𝑑 = where Cp is the heat capacity, is the material density, T is the temperature, k is the thermal conductivity, and q is the volumetric heat input. The FE thermal model uses a moving double ellipsoid Goldak heat source to simulate the arc power source [21]. The heat source includes two ellipsoidal sources that are combined to simulate the different temperature gradients at the front and rear of the melt pool.

The heat source model has been used in many modeling efforts and was created to simulate welding as well as metal AM [9], [11], [18], [20], [22], [23]. The double ellipsoidal parameters used for the FE thermal simulation can be found in Table 2.𝛿𝛿 2 𝑎𝑎𝑟𝑟𝛿𝛿𝛿𝛿22𝑞𝑞(𝑥𝑥, 𝑦𝑦, 𝑧𝑧, 𝑡𝑡) =− 3𝑏𝑏− 3𝑏𝑏�Table 2. The Goldak heat source parameters used for the FE thermal simulation.Factor Efficiency Half width (mm) Depth (mm) Front Length (mm) Rear Length (mm) Front Fraction (mm) Rear Fraction (mm)Symbol Valuen b c af ar ff fr0.95 3.25 3.25 7 13 0.7 1.3The boundary conditions applied to the thermal simulation include free heat convection and thermal radiation. The surface convection coefficient applied to the part and substrate is 18 W/(m2K) [24]. Radiation emissivity is held constant during the simulation being applied to the entire surface of the part and substrate with a value of 0.2 [22].

Radiation emissivity is held constant during the simulation being applied to the entire surface of the part and substrate with a value of 0.2 [22]. As the heat source travels along the print path, elements are progressively activated to simulate the deposited layer. Abaqus simulates the progressive addition of a material by modifying the value of its thermal conductivity: at the initial state, a very low value is assigned to the thermal conductivity of the elements prior to being activated, and then, the conductivity is increased to its actual value when the element temperature exceeds a threshold value [10]. Due to the low thermal conductivity of inactive elements, heat transfer is not able to occur1438between active and inactive elements –making it appropriate to simulate material deposition. Like the actual WAAM process, the temperature increases in the inactive elements only when they are directly heated by the Goldak heat source.

Like the actual WAAM process, the temperature increases in the inactive elements only when they are directly heated by the Goldak heat source. The mesh consists of 225,450 CD3D8 elements and can be seen in Figure 2.Figure 2. The FE mesh consisting of 225,450 linear brick elements (left) and the nodal extraction locations for layers 0,1,5, and 10 (right)..FE Thermal Simulations PerformedThree variations in input thermal properties were simulated for wire and as-built M250 material in the FE thermal simulations described previously. The variations include the experimentally measured values from DSC and LFA tests, linearized values for each property, and constant values at room temperature.

The variations include the experimentally measured values from DSC and LFA tests, linearized values for each property, and constant values at room temperature. The measured values were input as recorded for tests “wire_exp” and “waam_exp.” For the linearized properties, the first and last values were used to remove nonlinear portions for “wire_lin” and “waam_lin.” Lastly, constant values at room temperature were chosen for tests “wire_RT” and “waam_RT.” A description of each thermal input variation is shown in Table 3.Table 3.

The thermal property input details for the three test cases for each material.Naming Convention wire_exp waam_exp wire_lin waam_lin wire_RT waam_RTMaterial Type wire waam wire waam wire waamThermal Properties experimentally measured experimentally measured linearized properties linearized properities constant value at room temperature constant value at room termperatureM250 temperature dependent thermal properties were calculated for the weld wire and printed material by DSC and LFA via Netzsch and can be found in Figure 3. The properties for A36 were found from literature [25]. The values were measured up to a maximum temperature of 1250 °C with a heating rate of 293 °C/min.1439Figure 3. The measured and linearized thermal conductivity (left) and specific heat (right).There was a slight difference in density from 7.94 to 8.48 g/cm3 for wire and as-built material.

The measured and linearized thermal conductivity (left) and specific heat (right).There was a slight difference in density from 7.94 to 8.48 g/cm3 for wire and as-built material. The latent heat of fusion for M250 wire is 392.5 J/g with a solidus and liquidus temperature of 1443.5 °C and 1471.7 °C. The latent heat of fusion for the as-built M250 is 202 J/g with a solidus and liquidus temperature of 1450 °C and 1490 °C. A smoothing factor of 5 was utilized to improve convergence due to the sharp change in enthalpy in the simulation, which was also implemented in the Abaqus AM package example file [10].Results and DiscussionThe temperature history for a 10-layered thin wall built was simulated using the FE model described previously for each of the three thermal property variations for wire and as-built material. Nodal temperature was recorded at each of the thermocouple locations as well as along layers 0, 1, 5, and 10.

Nodal temperature was recorded at each of the thermocouple locations as well as along layers 0, 1, 5, and 10. The FE thermal history for layers 1-10 and 6-10 for thermocouple (TC0) are shown below in Figure 4.Figure 4. The FE thermal history at thermocouple TC0 for all layers (left) and layers 6-10 (right).The nodal temperature values at the location of TC0 appear to deviate based on the thermal property input used. Considering thermocouples located on the substrate are used for calibration of the heat source parameters and boundary conditions, it is important to understand if thermal property deviation can be detected in the temperature history at these locations. On the substrate, the data shows that the constant room temperature properties for wire (wire_RT) resulted in a decrease in temperature; however, for the as-built material (waam_RT) the opposite was true.

The linearized and experimental properties for the wire had the least change in temperature on the substrate compared to the other properties. For the as-built material properties, the linearized values resulted in the lowest thermal response while room temperature properties led to the highest temperatures.1440To contrast the simulated temperature variations on the substrate, nodal temperature data was investigated on the part surface for layers 0, 1, 5, and 10. If large thermal variation is seen in the part, the variation in heating and cooling cycles would most likely lead to unique stress profiles that may not be detected with measurements on the substrate. The simulated thermal history for nodes centered on the build located on the top surface of layers 0, 1, 5, and 10 is shown for each thermal input variation in Figure 5.Figure 5.

The simulated thermal history for nodes centered on the build located on the top surface of layers 0, 1, 5, and 10 is shown for each thermal input variation in Figure 5.Figure 5. The FE thermal history at each nodal location for six different thermal property inputs on layer 0 (a), layer 1 (b), layer 5 (c), and layer 10 (d).Like the temperatures measured at TC0, the values measured along several layers appear to show variation in thermal response over time for the printed layer, but the difference in temperature is much greater. The thermal history is shown to deviate substantially more in the part than along the substrate meaning the stress response would most likely deviate as well. The constant values at RT resulted in the largest temperatures for both wire and as-built material. Like the thermocouple data, the linearized and experimental values for wire material resulted in the least temperature variation of all groups tested.

Like the thermocouple data, the linearized and experimental values for wire material resulted in the least temperature variation of all groups tested. Overall, linearized values for both the wire and as- built material were shown to have reductions in temperature compared to the experimental and RT properties.To understand how the melt pool may be affected by the change in thermal properties, the maximum temperature during the simulation was recorded. The values for maximum temperature are artificially high due to the few number of elements provided in the beginning of each layer. The applied volumetric heat flux is being applied to a single row of elements leading to large spikes in maximum temperature shown on the left in Figure 6. It is believed that following the large spikes in temperature, the value reaches a steady state temperature before cooling during each layer. This is illustrated by the plateau seen following the large spike.

This is illustrated by the plateau seen following the large spike. To better demonstrate the steady state temperatures, the maximum temperature at the selected nodes along layers 0, 1, 5, and 10 is shown on the right in Figure 6.1441Figure 6. The maximum temperature across the entire simulation for every increment (left) and the maximum temperature recorded at each nodal location measured along the center of the part at layers 0, 1, 5, and 10 (right).The trend for thermal variation is similar as the data previously discussed for each of the thermal input properties. The properties held constant at RT for both the wire and as-built material appear to have the largest maximum temperatures, while the linear properties had the lowest maximum temperatures for each material type. The maximum values at each of the selected nodes can be found in Table 4.Table 4.

The trend described earlier is clear in table format. The constant values at RT for wire and as-built WAAM material were seen to have the highest maximum temperatures followed by experimentally measured properties, and linearized values.ConclusionFE thermal simulations of the WAAM process have been used to predict the temperature history of a 10 layered thin-wall using variations in temperature-dependent thermal properties: thermal conductivity, specific heat, and latent heat. Wire filler material and as-built WAAM printed material for M250 were collected experimentally using DSC and LFA to accurately simulate the thermal history and solidus to liquidus transition seen in the heat affected zone. The results shown suggest that the simulated thermal history can vary significantly depending on the material form selected (wire or as built), as well as the way in which the data is implemented into the thermal model.

Further work needs to be performed before determining the most appropriate source for thermal properties (wrought, wire, as-built, or material database), but it can be concluded that the form of material does affect the thermal response for maraging grade 250 steel. It was shown that constant values at room temperature result in an increased thermal response for WAAM thermal simulations, and linearized properties decreased the temperature when both were compared with the experimental thermal property inputs.AcknowledgementsThis work was supported by National Center for Manufacturing Sciences (NCMS) and US Army Research Laboratory (ARL); Collaboration Agreement #202031-141054 CTMA Cooperative Agreement #HQ0034-15-2- 0007.1442References[1] D. Ding, Z. Pan, D. Cuiuri, and H. Li, “Wire-feed additive manufacturing of metal components: technologies, developments and future interests,” Int J Adv Manuf Technol, vol. 81, no. 1–4, pp. 465–481, Oct.

FECTS OF AGING ON EPOXY-BASED RAPID TOOLING MATERIALSXavier Ottemer and Jonathan S. Colton Center for Polymer Processing, Rapid Prototyping and Manufacturing Institute, and School of Mechanical Engineering Georgia Institute of Technology, Atlanta, GA 30332-0405 USAAbstractThis study shows the effects of aging on epoxy-based rapid tooling materials. Two epoxy-acrylate resins used for stereolithography and one aluminum powder-filled epoxy used for the high speed machining of plastic injection molds were selected to investigate their aging behavior in different environments. Four different conditions, each characterized by a specific relative humidity, were used to perform a seven week long aging study. Temperature was kept constant at 30°C. Both stereolithographic resins showed a drop in mechanical properties and in their glass transition temperatures in wet environments, whereas aging time showed little influence on these properties.

Both stereolithographic resins showed a drop in mechanical properties and in their glass transition temperatures in wet environments, whereas aging time showed little influence on these properties. Moisture uptake is mainly responsible for that drop and coefficients of water diffusion were determined. On the other hand, the aluminum powder-filled epoxy showed little humidity-dependent drop in mechanical properties and in its glass transition temperature during aging. This result correlates well with the very low moisture absorption of this material.IntroductionPrototype tooling for injection molding applications needs to have a long lifetime regardless of storage conditions. Many RP’s materials are polymer-based. It is imperative therefore that test methods be developed in order to predict the long-term durability of these materials in different environments.

Many RP’s materials are polymer-based. It is imperative therefore that test methods be developed in order to predict the long-term durability of these materials in different environments. This study (Ottemer, 2001) focuses on three materials that can be used for soft tooling for injection molding applications: Somos SL7110, Ciba SL7510, and Vantico Renboard Ren Shape-Express 2000. Both SL7110 and SL7510 are epoxy-acrylate photopolymers for stereolithographic systems manufactured by 3D-Systems. Ren Shape-Express 2000 is an extremely fast milling aluminum powder-filled epoxy for prototype and short run tooling. These three materials are representative of the current trend in polymer-based rapid tooling materials.Aging of polymersAging is the term given to various processes that degrade polymers over time. During aging, one or more environmental stress factors act on the bulk polymer.

During aging, one or more environmental stress factors act on the bulk polymer. The effects of these stress factors are translated to the microscopic/molecular level where changes occur in the polymer network or chain. Allred (1981) used the glass transition temperature as an indicator of any physical changes in the resin due to aging.256Hydrolytic aging is the greatest concern as contact with water is almost unavoidable. Moisture has the most significant effect on the reduction in the properties of these materials.

Moisture has the most significant effect on the reduction in the properties of these materials. Moisture exposure has two important components:Diffusion of water into the sample: the moisture absorption of a resin can be attributed to the water's affinity for the highly polar functional groups in the cured resin.Plasticization of the resin: on a molecular level the effect of water associating with polar groups is to decrease the hydrogen bonding between polymer chains which causes plasticization of the resin (Roe, 1991).Time-humidity superposition principleMoisture accelerates the aging process in the same way as temperature.

A time-moisture principle has been proposed by Knauss and Kenner (1980) where the effect of moisture could be explained by the time-humidity superposition principle using a shift factor as shown in Equation 1.cB ( 1 Bc) 0 c− c −+log(ca ())−=(1) Where a(c) is the shift factor; B1, B2 are constants determined by regression; c, c0 are the moisture concentrations at time t, and at an arbitrary reference time, respectively.20Sample PreparationSamples made of SL7110 for tensile, diffusion, and DMA tests were built with 3D- Systems's stereolithography machine SLA250. Samples made of SL7510 were built with a SLA3500. Parts were all built vertically on the platform. After part formation, the excess resin was removed by rinsing with tripropylene glycol monomethylether, followed by a rinse in isopropyl alcohol. Then the parts were post-cured for 60 minutes with UV light and for 120 minutes in a thermal oven at 80ºC.

Then the parts were post-cured for 60 minutes with UV light and for 120 minutes in a thermal oven at 80ºC. RenBoard was shaped into samples using a conventional milling machine.Study of Moisture DiffusionTime dependence of coefficient of diffusionThe diffusion behavior of glassy polymers cannot be described adequately by a concentration-dependent form of Fick's law with constant boundary conditions. Deviations from Fickian behavior are considered to be associated with the finite rates at which the polymer structure changes in response to the sorption or desorption of penetrant molecules. This viscous behavior of the polymer makes the diffusion time-dependent. Three different equations were used to fit the data because the diffusion behavior of the resin systems did not strictly follow any well-known model. Coefficients of diffusion were determined by plotting the short-time, average, and long-time diffusion data using equations from Crank (1975).

Coefficients of diffusion were determined by plotting the short-time, average, and long-time diffusion data using equations from Crank (1975). For short times, Equation 2 was used.2571 2M M16tD s 2    t=Lπ(2) Where Mt is the mass uptake at time t, M∞ is the mass uptake at infinite time, Ds is the coefficient of short-term diffusion, t is the time, and L is the thickness of sample.

Coefficients of long-time diffusion were computed using Equation 3, where Dl is the long term diffusion coefficient.∞M Mt∞1 −=8 2 π − exp  2 tD π l 2 L 4  Then, assuming that the coefficient of diffusion is independent of solvent concentration, an estimated average of this coefficient can be derived with an error being about 0.001 percent using Equation 4 where the value of (t/L2) for which Mt/M∞ = 1/2 is defined as the half-time, (t/L2)1/2.D = av0.049 ( )1/2 2 t/LMoisture uptake measurements:ASTM D-570 is a standard test practice for water absorption. Eight standard specimens of each material were made. The dry weights of the samples were recorded with an accuracy of 0.1 mg. Then half of each kind of sample were put in pure water (humidity=100% and T=30ºC) and the remaining half in saturated salt-water. This aqueous solution has an equivalent relative humidity of 75.1±0.1% at 30°C, according to the ASTM E104.

This aqueous solution has an equivalent relative humidity of 75.1±0.1% at 30°C, according to the ASTM E104. However, this solution is not totally equivalent to a 75.1% RH environment because the osmotic pressure created by Na+ and Cl- ions will affect the leaching out of the polymer.The samples were weighed at 2 hrs, 4 hrs, 6 hrs (RenBoard), 8 hrs (RenBoard), 24 hrs, one week, and every two weeks until they absorbed less than 1% of the total weight gain or 5 mg per two weeks, whichever was greater. After all the test samples reached the termination criteria, they were placed in a dessicator overnight, and then reweighed in order to observe any leaching out of the materials. Figure 1 shows the weight gain versus time of the three materials. By comparing the main trends of the mass uptakes for the three materials, it appears that SL7510 absorbs the most water and RenBoard the least.

By comparing the main trends of the mass uptakes for the three materials, it appears that SL7510 absorbs the most water and RenBoard the least. Also, the infinite moisture absorption for both SL7110 and SL7510 is lower when immersed into salt-water, equivalent to a 75.1% RH environment, than in distilled water. That was not observed with RenBoard.The coefficients of diffusion are shown in Table I. These values are consistent with values of coefficients of diffusion of water in other commercial resins (Becker and Colton, 2000). RenBoard shows an odd behavior - the coefficients of diffusion for salt-water immersion are higher than those for water-immersion. This may be due to a chemical interaction between ions (Na+,Cl-) and the aluminum filler. Moreover a change of color from gray to brown has also been observed after 1 week of salt-water immersion and after 3 weeks of water-immersion.

Moreover a change of color from gray to brown has also been observed after 1 week of salt-water immersion and after 3 weeks of water-immersion. Only SL7510 immersed into water has shown some significant leaching out (3% of material loss).258(3)(4)76RenBoard-water RenBoard-salt-w54SL7510-water SL7510-salt-wat SL7110-waterSL7110-salt-wat32100100200300400500600700length of time of immersionFigure 1: Water uptakeTable I: Coefficients of diffusionmm2/s SL7510 - 100%RH SL7510 - 75%RH SL7110 - 100%RH SL7110 - 75%RH RenBoard - 100%RH RenBoard - 75%RHD short 10.0E-7 3.1E-7 3.3E-7 2.1E-7 2.1E-7 10.7E-7D average 21.8E-7 9.2E-7 9.4E-7 5.4E-7 8.5E-7 20.7E-7D long 13.8E-7 8.6E-7 2.3E-7 1.4E-7 2.5E-7 7.4E-7Aging studyEighty-four samples of each material were immersed in distilled water at 30°C. The same number of samples were put in a bath of saturated aqueous salt solution, equivalent to a 75.1% RH environment at 30°C.

The same number of samples were put in a bath of saturated aqueous salt solution, equivalent to a 75.1% RH environment at 30°C. Another set of samples were exposed to ambient atmosphere kept at 30°C (dry-bulb temperature) by the use of a ventilated thermal oven. The RH of the environment varied other time due to weather conditions. It was 30±5% at 30ºC during the aging study on SL7510, 23±5% at 30ºC during the study of SL7110, and 20±6% at 30ºC during the study of RenBoard.

It was 30±5% at 30ºC during the aging study on SL7510, 23±5% at 30ºC during the study of SL7110, and 20±6% at 30ºC during the study of RenBoard. Finally, a dessicator placed in the oven provided dry conditions for the last set of samples.Mechanical properties:259At specified aging times, two sets of three samples of each material were removed anddynamical mechanical analysis (DMA) and tensile test were performed.Storage modulus E’ and Glass Transition Temperature Tg.These characterizations were made with DMA testing (DMA 2980, TA Instruments, New Castle, DE) in the single cantilever flexural mode at a frequency of 1Hz and an amplitude of deformation of 20 _m. E’ and Tg were recorded during a temperature ramp of 3°C/min from 30°C to 130°C for both SL7110 and SL7510 and from 30°C to 270°C for RenBoard. Tg was measured at the peak of tan_=E”/E’.Young’s modulus and ultimate tensile propertiesStandard ASTM D638 was used as a guideline for the mechanical testing procedure.

Tg was measured at the peak of tan_=E”/E’.Young’s modulus and ultimate tensile propertiesStandard ASTM D638 was used as a guideline for the mechanical testing procedure. A tensile testing machine (Instron 4466, Canton MA) with a 10 kN load cell was used, and an extensometer was mounted on the sample. The cross head speed was 0.25 cm/min. Both strain and load were recorded.Aging behavior of SL7510 and SL7110:The evolution of the glass transition temperature is a good indicator of the aging behavior of a material. Figure 2 shows the evolution of the Tg of SL7510. After an induction period of around 400 hr due to diffusion of water inside the materials (area I on Figure 2), the Tg becomes time-independent and shows a marked decrease with RH due to the effect of moisture uptake on the material (area II on Figure 2). From 96ºC for the dry environment, the Tg of SL7510 drops down to 55ºC for the wet environment.

From 96ºC for the dry environment, the Tg of SL7510 drops down to 55ºC for the wet environment. This drop is an indication that the resin is being plasticized by moisture uptake.The results of the mechanical tests are similar for both SL resins, showing a drop in properties in wet environments, except for the yield strain which remains mostly unaffected and constant, whatever the aging time or the RH of the environment.After a 400 hr induction period, aged SL7510 samples show a 66% decrease in ultimate tensile strength from dry to wet environments. Initial properties were UTS = 54.5 MPa, E = 2407 MPa, E' = 2200 MPa, and ultimate strain = 5.75%. On the other hand, the tensile strength remains constant for the dry environment, with a slight increase observed. The storage modulus, E’, and the Young’s modulus, E, of SL7510 also reveal the same behavior (Figure 3): first an induction period followed by a time-independent, RH-dependent stage.

The storage modulus, E’, and the Young’s modulus, E, of SL7510 also reveal the same behavior (Figure 3): first an induction period followed by a time-independent, RH-dependent stage. It also is observed that the ultimate and yield strengths of SL7510 decrease with RH in the same way as Tg, E’, and E.260Area IArea II11010090) C (g T80700% RH 30% RH 75% RH 100% RH6050400200400600800100012001400aging time (hrsFigure 2: Tg of SL751035000%3000250045%200075%15001000100%50000200400600800100012001400aging timeFigure 3: Young's modulus of SL7510The same overall behavior is observed for SL7110, but to a lesser extent. Its initial values were UTS = 52 MPa, E = 2450 MPa, E' = 2000 MPa, and ultimate tensile strain = 4.6%. After an induction period of around 400 hr, the Tg becomes more time-independent and shows a marked decrease with RH (Figure 4).

After an induction period of around 400 hr, the Tg becomes more time-independent and shows a marked decrease with RH (Figure 4). From 108ºC for the dry environment, the Tg drops down to 75ºC for the wet environment. The mechanical test results indicate the same, showing a drop in properties in wet environments, but here the properties are ruled by a 200 hr induction stage. The ultimate tensile strength is 25.3% higher than initially in dry and ambient environments, remains constant when immersed into the 75.1% RH environment, and drops 12% from its unaged value when submitted to the wet environment.2611101009080700% RH605023% RH 75% RH 100% RH400200400600800100012001400aging time (hFigure 4: Tg of SL7110Aging behavior of RenBoard:RenBoard does not behave like the two other materials. The initial values were UTS = 52 MPa, E = 15000 MPa, E' = 6026 MPa, and the ultimate tensile strain = 0.83%.

The initial values were UTS = 52 MPa, E = 15000 MPa, E' = 6026 MPa, and the ultimate tensile strain = 0.83%. Figure 5 shows the evolution of the Tg of RenBoard with aging time. The variations observed are not significant. The Tg remains quite constant at 248ºC whatever the aging time or the RH of the environment. The mechanical results indicate the same, showing almost no significant variations. Whatever the aging time or the RH, the material failed in a brittle manner. This is different from the behavior of SL7510 and SL7110.250.0249.5249.0248.5248.0247.5247.0246.5246.0245.50% RH 20% RH 75% RH 100% RH245.0020040060080010001200aging time (hFigure 5: Tg of Renboard262ConclusionsThe aging behaviors of three epoxy-based rapid tooling materials were investigated under four different moisture environments at 30ºC over seven weeks.

Two were stereolithographic resins and showed a drop in mechanical properties and in their glass transition temperatures in wet environments, whereas aging time had no significant influence on these properties. The magnitude of the drop of mechanical properties was found to be related directly to the magnitude of the moisture uptake by the material. The properties (Tg and mechanical properties) first undergo an induction period led by moisture absorption and then reach asymptotic values, which are functions of the infinite mass-uptakes. The higher the moisture absorption, the more affected are the properties. The third material, an aluminum-filled epoxy, did not show significant humidity-dependent drops in mechanical properties and in its glass transition temperature.

The third material, an aluminum-filled epoxy, did not show significant humidity-dependent drops in mechanical properties and in its glass transition temperature. This result correlates well with the very low moisture absorption ability found for this material.AcknowledgmentsThis work has been made possible thanks to the National Science Foundation (DMI- 9618039) and the Rapid Prototyping and Manufacturing Institute, whose member companies include 3D-Systems, 3M, Baxter, Ciba, Club Car, Durden Enterprises, Ford, Kodak, Lucent Technologies, Pratt&Whitney, and Siemens.ReferencesAllred, R.E. (1981), Journal of Composite Materials, Vol. 15, No. 2, 117-132. Becker, K. and Colton, J.S. (2000), Polymer-Plastics Technology and Engineering, Vol. 39, No.4, 667-682.Crank, J. (1975), The Mathematics of Diffusion, 2nd Edition, Oxford: Oxford University Press. Knauss, W.G., and Kenner, V.H. (1980), Journal of Applied Physics, Vol. 51, No.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffects of Area Fraction and Part Spacing on Degradation of 304L Stainless Steel Powder in Selective Laser MeltingCaitlin S. Kriewall1, Austin T. Sutton2, Sreekar Karnati2, Joseph W. Newkirk1, Ming C. Leu21Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409 2Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractIn selective laser melting (SLM) systems, a large portion of powder remains unconsolidated and therefore recycling powder could make SLM more economical. Currently, a lack of literature exists specifically targeted at studying the reusability of powder. Furthermore, the definition of powder reusability is complex since powder degradation depends on many factors.

Furthermore, the definition of powder reusability is complex since powder degradation depends on many factors. The goal of the current research is to investigate the effects of area fraction and part spacing on the degradation of 304L powder in SLM. An experimental study was conducted where various area fractions and part distances were chosen and powder characterization techniques for determination of particle size distributions, tap and apparent densities, and x-ray diffraction were employed to track evolving powder properties for the purpose of reuse. The results show that the recyclability of 304L powder depends on the utilization of the build area causing varying degrees of particle size coarsening and delta ferrite formation.IntroductionSelective laser melting (SLM) is a powder bed fusion additive manufacturing (AM) technique where parts are built layer by layer. In this process, powder is dispensed and spread over a build plate at a predetermined thickness.

In this process, powder is dispensed and spread over a build plate at a predetermined thickness. A laser is then scanned over locations specified by a CAD model to consolidate the powder. After completion of the layer, the build plate is incremented down by one layer thickness and the process is repeated until the final part dimensions are achieved [1]. AM has attracted the interest of several industries due to its ability to make near net shaped parts with complicated geometry that are not attainable with subtractive processing technologies. However, powder bed processes inherently have a significant percentage of powder that is leftover after processing, as generally only 10%-50% of the build area is used for parts. Therefore, there is an opportunity to reuse this leftover powder, which would benefit the economy of the process as currently 5-46% of the cost of the SLM process is attributed to material costs [2].

Therefore, there is an opportunity to reuse this leftover powder, which would benefit the economy of the process as currently 5-46% of the cost of the SLM process is attributed to material costs [2]. However, powder characteristics and their effect on part properties must be well understood before widespread reuse of powder can be implemented.Although sparse, there are some studies in the literature on the recyclability of certain metal powders including Ti64, Inconel 718, 17-4 PH stainless steel, and CoCr [3-5]. Tang et al. [3] studied the recyclability of Ti64 by creating a build of tensile specimens in an Arcam EBM system and collecting the powder after each build. The powder composition, size distribution, apparent and tap densities, flowability and morphology were examined after selected uses. They found that277the powder could be reused up to 21 times with no undesirable effects on the mechanical properties. Ardila et al.

They found that277the powder could be reused up to 21 times with no undesirable effects on the mechanical properties. Ardila et al. [4] studied the reusability of Inconel 718 on a Model Realizer SLM 250. Powder was collected after each iteration and the particle size distribution and composition was studied. Parts were also produced in order to evaluate the porosity, toughness, hardness and microstructure. It was found that the Inconel 718 powder could be reused up to 14 times with no significant effect on the part properties. Slotwinski et al. [5] studied the reusability of 17-4 PH stainless steel and CoCr powders in the SLM process. Although investigation of the mechanical properties was not discussed in this article, it was shown that reusing the powders corresponded with an increase of the particle size distribution.

Although investigation of the mechanical properties was not discussed in this article, it was shown that reusing the powders corresponded with an increase of the particle size distribution. In all of these studies, the methodology behind the build design was not discussed; it was only presented as a certain build that was produced and iterated multiple times.Heat-affected powders are created during the SLM process and can deposit in the build area [6]. The area of parts has an effect on the amount of heat-affected powders that are produced. As this powder has been shown to have a deleterious effect on part properties [6], the success of a recyclability study will depend on the amount of parts that are present in the build. The effect of the area fraction and part spacing of built parts on the powder properties was investigated in this study.

The effect of the area fraction and part spacing of built parts on the powder properties was investigated in this study. This approach aims to be a simple way to design a recyclability study, as each of the builds is relatively small and would take a minimal amount of powder and machine time. This preliminary study can be used to assess the effect of different area fractions on powder degradation. Then, a build can be designed that takes into account the desired specimens and area fractions where powder changes are known to occur. In this way there is a better chance of actually capturing changes in part properties. This study could also aid in pinpointing the area fractions where there begins to be an effect on the part properties, if such a dependence is present in the material.Experimental MethodsGas-atomized 304L SS powder was purchased from LPW Technology with the chemical composition listed in Table 1.

Prior to processing, the powder was passed through a 63 µm mesh sieve in order to breakup agglomerates and remove large particles that would have a negative effect on the layer thickness and uniformity. The powder was used to build parts on a Renishaw AM 250 SLM machine which contained a pulsed Nd-YAG laser with a Gaussian profile beam intensity and a wavelength of 1070 nm. The diameter of the beam spot on the powder bed was approximately 70 µm. The oxygen content in the build chamber was kept below 1000 ppm and the substrate temperature was held at 80°C.

The diameter of the beam spot on the powder bed was approximately 70 µm. The oxygen content in the build chamber was kept below 1000 ppm and the substrate temperature was held at 80°C. Finally, a constant volumetric argon gas flow of 400 ft3/min was maintained across the build area during processing and was recirculated through the AM 250.Table 1: Chemical Composition of 304L SS powder used as the starting material in this studyElementCCrCuFe MnNNiOPSWt%0.018 18.4 < 0.1 Bal1.40.069.80.02 0.012 0.005278Si0.6Several builds were used in order to study the effect of area fraction and part spacing on the degradation of powder. Four of these builds consisted of 5x5 arrays of squares that were built at different area fractions, i.e. 12.5%, 25%, 50%, and 75% of the build chamber consisted of parts. This corresponded to squares with edge dimensions of 15 mm, 21 mm, 30 mm, and 36 mm for the 12.5%, 25%, 50%, and 75% builds.

12.5%, 25%, 50%, and 75% of the build chamber consisted of parts. This corresponded to squares with edge dimensions of 15 mm, 21 mm, 30 mm, and 36 mm for the 12.5%, 25%, 50%, and 75% builds. An important note is that the area fraction corresponds to the area enclosed by the 4 bolt holes as corners and not the entire base plate. An additional three builds were conducted in order to evaluate the part spacing on the degradation of the powder. These builds used a 50% area fraction with different arrays of squares built, i.e. 7x7, 9x9, and 11x11 with part spacings of 8 mm, 6 mm, and 5 mm respectively. These builds are illustrated in Figure 1. Each of the seven builds was replicated three times in order to study how the powder characteristics change with recycling. Powder was collected after each build from the overflow and build area, characterized, sieved, and then placed back into the AM 250 for the 2nd and 3rd iteration. No new powder was added for successive builds.

No new powder was added for successive builds. For the first iteration of builds, the squares were built up to 5 mm tall meaning that the builds ranged in time from 4-24 hours. For each subsequent iteration, the height of the parts was decreased by 1.5 mm.Figure 1: Illustration of the experimental builds. Note the 50% 5x5 build has been shown twice but was only completed once for each iteration. Part spacing builds were all at 50% area fraction.As this work is designed to be a relatively quick and easy way to determine builds for a more vigorous powder recycling study, a limited selection of particle characterization techniques279were used and they were chosen based on the usefulness of the information and/or the ease of the measurement. The powder characterization included investigation of the particle size distributions, tap and apparent densities, and the phases present in the powder leftover from each build.

The powder characterization included investigation of the particle size distributions, tap and apparent densities, and the phases present in the powder leftover from each build. Prior to any characterization, powder was mixed using a Turbula T2C mixer for a minimum of 15 minutes to ensure homogenization before sampling. Particle size distributions were obtained using an ASPEX 1020 scanning electron microscope (SEM) using the automated feature analysis (AFA) capabilities of the instrument. Each particle size distribution was generated from at least 2000 powder particles in order to establish an accurate distribution [7]. The particle size distributions were fit using JMP Pro software to obtain a 95% confidence interval of the distribution. Measurements of tap and apparent densities were conducted following ASTM standards B527-15 and B329-14 respectively. Each measurement of tap and apparent density was replicated three times.

Measurements of tap and apparent densities were conducted following ASTM standards B527-15 and B329-14 respectively. Each measurement of tap and apparent density was replicated three times. Determination of the percentage of phases was conducted using a Panalytical X’Pert Pro Multi-Purpose Diffractometer X-Ray Diffraction instrument. Due to the time and cost of the instrument, XRD scans were only obtained for samples of the 12.5%, 50% 5x5, 50% 11x11, and 75% builds in order to sample the extremes. XRD scans were performed from a 2θ range of 20- 100 over a 3 hour period. RIQAS 4 software was used to perform the Rietveld refinement for quantification of phases.

XRD scans were performed from a 2θ range of 20- 100 over a 3 hour period. RIQAS 4 software was used to perform the Rietveld refinement for quantification of phases. Finally, the powder after each build was sampled for testing and sieved through a 63 µm mesh sieve before the next build and the amount of powder sieved off was documented.Results and DiscussionParticle size distributions have significant effects on the processing behavior of metal powders and have been shown to have an effect on the built part properties [8, 9]. The D10, D50, and D90 values from the numeric particle size distribution fit using JMP Pro are shown in Table 2. Here the plus or minus values show the 95% confidence interval. Each D-value corresponds to powder that is less than or equal to a certain size at the specified percentage. For example, the first iteration of the 12.5% area fraction build has a D50 value of 22.3 µm.

For example, the first iteration of the 12.5% area fraction build has a D50 value of 22.3 µm. This means that in the cumulative number plot, 50% of the powder was less than or equal to 22.3 µm. In order to make conclusions about the particle size distributions, the three D values must be considered together.When considering the area fraction builds, both the D50 and D90 (shown in Figure 2 and Figure 3 respectively) values showed a statistically significant increase of the particle size between iteration 1 and 2. This increase was not observed for the D10 values, where for most cases (excluding the 50% 5x5), the D10 either was not statistically significantly different or there was a decrease observed. This coupled with the increase in D50 and D90 for the area fraction builds indicated that the particle size distributions were spreading out between iteration 1 and 2.

This coupled with the increase in D50 and D90 for the area fraction builds indicated that the particle size distributions were spreading out between iteration 1 and 2. However, this behavior was not carried into iteration 3, where the only area fraction build with a statistically significant increase in particle size distribution was the 25% build. The remaining area fraction builds showed a decrease in D90 values and either a decrease or no change in the D50 values. A trend was unable to be established for the D10 values between iteration 2 and 3. The coarsening observed between iteration 1 and 2 in the area fraction samples was not observed reliably in the part spacing samples. The 50% 5x5 build and the associated trends were discussed with regards to280the area fraction.

Error bars show the 95% confidence intervals.281Iteration 335.4 ± 0.636.4 ± 0.634.2 ± 0.536.1 ± 0.638.2 ± 0.835.4 ± 0.537.5 ± 0.6464442m µ40, e z i S e l c i t r a P383634Iteration 1Iteration 2Iteration 33230282612.5% 5x525% 5x550% 5x575% 5x5Figure 3: D90 values obtained from numeric particle size distributions for each iteration for the seven builds. Error bars show the 95% confidence intervals.Table 3 shows the apparent density, tap density, and Hausner ratio of powder samples for each iteration, where the apparent and tap densities are graphically displayed in Figure 4 and Figure 5. Both the tap and apparent densities increased between iteration 1 and iteration 2 for all of the builds. Tap densities also increased between iteration 2 and iteration 3 in all cases except the 50% 9x9 build. However, the apparent densities dropped after iteration 2.

Tap densities also increased between iteration 2 and iteration 3 in all cases except the 50% 9x9 build. However, the apparent densities dropped after iteration 2. With the increase of both the tap and apparent densities at iteration 2, a decrease in the Hausner ratio for certain builds (50% and 75%) was observed. The Hausner ratio is the ratio of the tap to apparent densities and serves as a simple way to describe the flowability of powders [10]. Although this ratio has been noted to be unsatisfactory in completely describing the flowability in AM processes, the decrease in Hausner ratio corresponds with improving powder flowability. Improvement of powder flowability with reuse has also been noted by other researchers using different materials [3]. This improved flowability could imply that the ideal powder may not be the virgin powder.

This improved flowability could imply that the ideal powder may not be the virgin powder. As both the tap and apparent density values were increasing, the density of each powder layer and therefore the bed density was expected to improve as well [9]. This improved bed density could have a positive influence on the part density, as a more dense powder bed will increase the density of the part by decreasing the lack-of-fusion pores.

Error bars show the standard deviations of the three iterations.2835.2Area FractionPart Spacing5.1L m / g , y t i s n e D p a T5.04.94.84.74.6Virgin - 63 µm12.5% 5x525% 5x5 50% 5x5 75% 5x550% 5x5 50% 7x7 50% 9x950% 11x11Figure 5: Tap densities for various builds after each iteration. Error bars show the standard deviations of the three iterations.The results of the Rietveld refinement XRD pattern fitting on the 12.5%, 75%, 50% 5x5, and 50% 11x11 builds are shown in Table 4. In this table, the weight percentage of δ-ferrite is shown as each sample only had two identifiable phases, δ-ferrite and γ-austenite. 304L SS is an austenitic stainless steel that, according to the chrome nickel equivalent, follows a ferritic- austenitic solidification path where:L → L + δ → L + γ + δ → γ + δ → γwhere δ is δ-ferrite and γ is γ-austenite [11].

Fast cooling is necessary to retain δ-ferrite [12] however at extremely high cooling rates (on the order of 105 – 106 K/s), the supercooled liquid can be present below the metastable solidus and solidification by primary austenite is possible [13]. Therefore, in the 304L SS powder processed through SLM, δ-ferrite becomes a marker of heat- affected powder, as the virgin powder has very little δ-ferrite due to the extremely high cooling rates achieved through the gas atomization process. In contrast, various kinds of heat-affected powder are present in the process due to the various interactions with the laser that have cooling rates sufficient to retain δ-ferrite [6]. For comparative purposes, the amount of δ-ferrite in the virgin 304L SS powder that had been passed through the 63 µm sieve was 1.5%.When considering the area fraction builds (12.5% and 75%), there was an obvious difference in the amount of δ-ferrite present in the powder samples.

In each iteration, the amount of δ-ferrite in the 75% build was always higher than in the 12.5% build. Both builds showed an increase, however the 75% build experienced a greater increase in δ-ferrite after each iteration. The part spacing builds are again more difficult to interpret, as there seems to be an anomalous result in the 50% 5x5 iteration 1. Comparisons between 50% 5x5 and the 11x11 builds do show that the 11x11 build has more δ-ferrite. Additionally, the amount of δ-ferrite increases with each iteration (for the 11x11 builds) and appears to increase at a faster rate than the 5x5 builds. These284Iteration 1Iteration 2Iteration 3results showed that both the area fraction and the part spacing have an effect on the amount of heat-affected powder that is generated and deposited inside the build chamber during processing.Table 4: Rietveld refinement results for quantification of XRD spectra.

Values shown are weight percent of δ-ferriteArea FractionArray SizeIteration 1Iteration 2Iteration 312.5%5x50%0.1%1.3%50%5x52.2%0.9%1.1%50%11x114.2%4.9%6.1%75%5x53.6%7.2%11.4%Figure 6 shows the results of the percentage of powder that was sieved off prior to being run for the next iteration. Between iteration 1 and 2, more powder was generally sieved off, although the amount sieved off for the 12.5% build stayed consistent. Furthermore, between iteration 2 and 3 the majority of the samples either stayed the same or decreased in the amount of powder sieved off, although the 50% 5x5 and 9x9 builds still increased. This result was consistent with what was observed from the particle size distributions. When considering the different area fractions, there was a clear trend that increasing the area fraction corresponds to an increase in the amount of powder that was sieved out.

When considering the different area fractions, there was a clear trend that increasing the area fraction corresponds to an increase in the amount of powder that was sieved out. Again, it was more difficult to establish a trend concerning the part spacing builds.12.0%g n i v e i S m o r f10.0%8.0%Area FractionPart Spacings s o L r e d w o P f o t n e c r e P6.0%4.0%2.0%0.0%12.5% 5x525% 5x5 50% 5x5 75% 5x550% 5x5 50% 7x7 50% 9x950% 11x11Figure 6: Percentage of powder sieved off after each iteration of builds.Powder from any of these builds has been shown to change in at least one of the characterizations performed. However, the degree to which it changes was different for each build.

However, the degree to which it changes was different for each build. For the 12.5% area fraction builds, the change was noted in the particle size distributions and tap and apparent densities, yet the XRD and percent of powder sieved off show fairly consistent285Iteration 1Iteration 2Iteration 3powder with minor differences (where the amount sieved off increased from 0.8% to 0.9% and the weight percent of δ-ferrite increased by only 1.3%). On the other hand, the 75% powder showed differences in every test, however a 75% area fraction build would be difficult to justify for a recycling study as the first build would have to be repeated several times in order to generate enough powder to carry through the study, depending on how many iterations would be needed. A 304L SS powder reusability study incorporating more mechanical testing and powder characterization parts will be employed.

A 304L SS powder reusability study incorporating more mechanical testing and powder characterization parts will be employed. Area fractions between 25% and 50% will be used for this study, because powder in this range has detectable differences that have been shown by other researchers to have an effect on the part properties.Summary and RemarksPowder characterization of the area fraction builds (12.5%, 25%, 50%, and 75%) and part spacing builds (50% 5x5, 7x7, 9x9, and 11x11) that were replicated three times enabled insights into the effect of area fraction on the powder. These included that increasing the area fraction induced an increase in the D50 and D90 of the particle size distributions between iteration 1 and 2. The corresponding D10 values stayed consistent or lowered, meaning that the particle size distributions were not only coarsening, but also widening.

The corresponding D10 values stayed consistent or lowered, meaning that the particle size distributions were not only coarsening, but also widening. However, this behavior was not carried into iteration 3, where particle size distributions were shown to drop, yet the D50 and D90 were still higher for iteration 3 compared to iteration 1. Tap and apparent densities both increased between iteration 1 and 2 for all of the area fraction builds. This improvement in tap and apparent densities coupled with the spreading of the powder could have an advantageous effect as the powder bed density was expected to increase. While the tap density increased further on iteration 3, the apparent density either remained consistent or fell to a lower value for this iteration. The XRD analysis showed that the 75% area fraction had more δ-ferrite, an indication of the amount of heat-affected powder, than both the 12.5% and 50% area fraction builds.

The XRD analysis showed that the 75% area fraction had more δ-ferrite, an indication of the amount of heat-affected powder, than both the 12.5% and 50% area fraction builds. Finally, the amount of powder sieved off increased with increasing area fraction for all iterations. When comparing the same build through different iterations, the amount of powder sieved off was shown to increase between iteration 1 and 2 apart from the 12.5% area fraction that stayed consistent. Furthermore, between iteration 2 and 3 the powder sieved off was not as consistent, where 12.5% build stayed consistent, 25% and 75% decreased, and 50% increased. The amount of powder sieved was shown to change significantly for the 75% build with each iteration however the 12.5% build stayed consistent through each iteration.Across all the powder characterization techniques employed it was more difficult to establish trends with the part spacing builds.

Trends were not identifiable in the particle size distribution data. The same trends observed in the area fraction builds concerning the tap and apparent densities were observed in the part spacing builds. The XRD showed that the 11x11 builds had more δ-ferrite compared to the 5x5 build and the amount of δ-ferrite in the 11x11 build increased with each iteration. Finally, the sieving results showed no trend when comparing the different part spacing builds, however, when comparing iteration 1 to 2 there was an increase in the powder sieved off for each part spacing build. The 5x5 and 9x9 showed an increase in the286amount of powder sieved off between iteration 2 and 3 while builds 7x7 and 11x11 had a decrease in the amount of powder sieved off.Powder from each of the builds was shown to have differences in at least one of the characterizations performed, however the degree to which the powder changed was different.

The 75% build showed the most drastic differences, however this area fraction would be unrealistic for a more involved reusability study. For the more involved 304L SS powder reusability study, area fractions between 25% and 50% will be used given that the powders in this range did show detectable differences that could have an effect on part properties. Future work will involve powder characterization on samples post-sieving to correlate the heat-affected powder to the sieve loss. Additionally, mechanical testing will be needed to evaluate whether the differences observed in the powder equate to differences in part performance. For this testing, mini-tensile specimens will be extracted at 5 locations corresponding to the diagonal across the build plate. Comparisons between each build and iteration and the different builds will be completed. Ideally, this work will provide a way to optimize the powder loss to machine productivity trade-off.

Comparisons between each build and iteration and the different builds will be completed. Ideally, this work will provide a way to optimize the powder loss to machine productivity trade-off. Finally, a more involved 304L SS powder reusability study will be performed.AcknowledgmentsThis work was funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.Furthermore, the authors would like to acknowledge Jeff Hill, Jacob Ficht, and EricBohannan for their assistance in this research.References[1] A. T. Sutton, C. S. Kriewall, M. C. Leu, and J. W.

Effects of build orientation and heat treatment on neutron-induced hardening were investigated by inspecting L-PBF samples built vertically or at a 45º angle in the following heat-treated conditions: as-built (no heat treatment), 700 ºC for 1 hour, 900 ºC for 1 hour, and 1050 ºC for 1 hour. The microhardness results of L-PBF samples before and after neutron irradiation were compared with traditional wrought Inconel 625. All samples underwent an irradiation flux of 6.61 x 1013 neutrons/cm²/s for 310 hours for an estimated damage of 0.012 dpa. Results indicate that as-built L-PBF specimens are less prone to radiation hardening relative to their wrought counterparts.

Results indicate that as-built L-PBF specimens are less prone to radiation hardening relative to their wrought counterparts. As-printed diagonal specimens were shown to harden by 8% as compared to 1.2% hardening in as-printed vertical specimens.Keywords: Selective Laser Melting, Nickel Superalloy, Nuclear Radiation Damage, Vickers Microhardness, Additive ManufacturingIntroductionIn recent decades, researchers have been working on building smaller, simpler, and more efficient nuclear reactors like Small Modular Reactors (SMR) and GEN IV reactors. Demand is high for structural materials used in building in-core and out-core components of GEN IV reactors that can withstand high operating temperatures and radiation doses to improve the life and safety of the reactor. Additive Manufacturing (AM) serves as a novel method to build complex-shaped components layer by layer with designed mechanical properties for higher temperatures.

Additive Manufacturing (AM) serves as a novel method to build complex-shaped components layer by layer with designed mechanical properties for higher temperatures. AM provides advantages compared to conventional machining like design flexibility, faster process setup, cost and waste reduction, and real time monitoring of microstructure and mechanical properties. A relatively mature and common technique for the AM of metals is laser-powder bed fusion (L-PBF) which employs 20-60 micron-thick layers of powder that is selectively melted1026using a laser beam within an inert atmosphere. For L-PBF, un-melted powder can help support new layers, thus allowing for the manufacture of complex components like heat exchangers, cellular structures, and topology-optimized parts.Nickel based superalloys are proposed to be utilized in GEN IV reactors [1]. One nickel- based superalloy, Inconel 625, possesses excellent high temperature strength, creep, and toughness.

One nickel- based superalloy, Inconel 625, possesses excellent high temperature strength, creep, and toughness. It is gaining significant importance to be utilized in next-generation nuclear reactors particularly in control rod and reactor core applications [2]. Inconel 625 is a solid solution strengthening alloy due to the presence of elements like Mo and Nb. Inconel 625 is becoming more ubiquitous in various industries like aerospace, nuclear or any other industries where harsh environments are expected [3]. Conventional machining methods often face difficulties in processing Inconel 625 components due to its high hardness, reduced machinability, and low thermal conductivity [4]. There have been various studies conducted in the past to understand the microstructure and mechanical properties of Inconel 625 built via different AM techniques. Fang et al. [5] studied the microstructure evolution of Inconel 625 components built via L-PBF that were subjected to different heat treatments.

Fang et al. [5] studied the microstructure evolution of Inconel 625 components built via L-PBF that were subjected to different heat treatments. It was found that as-built specimens possessed higher hardness due to high dislocation density arising from high thermal gradients during the L-PBF process. The microhardness values decreased when the as-built specimens were subjected to heat treatment temperatures at 870, 980, and 1150 oC for 1 hour. The samples were furnace cooled to room temperature. The heat treatment reduced the dislocation density and increased homogenization of grains. Szmytka et al. [6] compared the mechanical properties of Inconel 625 fabricated via L-PBF and Laser Melting Deposition (LMD) – a Directed Energy Deposition (DED) AM method, as well as wrought. They found that as-built L-PBF and LMD specimens exhibited higher hardness values compared to wrought specimens.

They found that as-built L-PBF and LMD specimens exhibited higher hardness values compared to wrought specimens. In addition, they discovered after performing a 1h heat treatment at 1100 oC, the hardness values decreased and thus the ductility of the specimens increased.Neutron irradiation on metals induces defects like voids, interstitials, dislocations, swelling, helium embrittlement, and radiation induced hardening [7][8]. While there have been studies conducted in the past to investigate the influence of ion irradiation on AM Inconel metals, the effects of neutron irradiation on mechanical properties of AM Inconel 625 are still unclear [9][10]. Claudson studied the effects of neutron irradiation, with exposures exceeding >1 MeV, on the elevated temperature mechanical properties of wrought nickel-based and refractory metal alloys.

Claudson studied the effects of neutron irradiation, with exposures exceeding >1 MeV, on the elevated temperature mechanical properties of wrought nickel-based and refractory metal alloys. The results showed a decrease in ductility at elevated temperatures compared to those measured at room temperature [11].The aim of this paper is to investigate the effects of neutron irradiation on the hardness of wrought and AM Inconel 625. The AM Inconel 625 samples were built using two different build orientations (vertical and diagonal) via L-PBF. All L-PBF and wrought Inconel 625 samples were irradiated inside a 10 MW nuclear reactor for duration of 310 hours. During this irradiation, each sample experienced a neutron flux of 6.61 x 1013 neutrons/cm2/s. Out of this total flux, 90% was contributed by thermal neutrons and 10% by fast neutrons. The total flux corresponds to a displacements per atom (dpa) value of 0.012.

Out of this total flux, 90% was contributed by thermal neutrons and 10% by fast neutrons. The total flux corresponds to a displacements per atom (dpa) value of 0.012. The radiation induced hardening was measured using Vickers microhardness on all pre- and post-irradiated samples. Comparing the hardening behavior of additive manufactured components and wrought Inconel 625 under neutron irradiation provides1027a better understanding on how AM and post processing affect the structural integrity of an important nickel-based superalloy under harsh environment.MethodsPlasma atomized, virgin Inconel 625 powder with a mean particle diameter of 31.31 µm was supplied by Carpenter Technologies. The chemical composition of Inconel 625 powder used for L-PBF in this study, adhering to ASTM Standard for Additively Manufacturing Nickel Alloy UNS N06625, is presented in Table 1. Inconel 625 cubes of 10 mm3 dimension were additively manufactured using a Concept Laser Mlab 100R system.

Inconel 625 cubes of 10 mm3 dimension were additively manufactured using a Concept Laser Mlab 100R system. The Concept Laser consists of a 100 W fiber laser with wavelength of 70 nm. The building process was carried out in an enclosed argon atmosphere for minimization of oxidization. Cubes were iteratively built at various process parameters and then inspected for density. Final process parameters provided for 99.79% dense samples, these process parameters were: power = 90 W, scan speed = 800 mm/s, laser diameter = 80 µm, hatch spacing = 60 µm, and a layer thickness = 25 µm. For experimental control and baselining, wrought Inconel 625 plate (152.4 x 152.4 x 4.76 mm3) was procured from Metalmen, NY, USA. The plate was manufactured per AMS5599 specification. Small 1 x 1 x 0.5 mm3 samples were sectioned from the plate using wire electrical discharge machining (EDM).

The plate was manufactured per AMS5599 specification. Small 1 x 1 x 0.5 mm3 samples were sectioned from the plate using wire electrical discharge machining (EDM). The chemical composition of wrought Inconel 625 is indicated in Table 1.Table 1: Elemental composition of L-PBF powder and wrought Inconel 625 (% in weight)% wt.UNS N06625L-PBFNi CrBal 20-23Bal 21.59Mo8-109.0Fe<5.02.95Nb+ Ta 3.15 - 4.15 3.55 0.03 0.04CoCPSAlTi Mn<1<0.1 <0.015 <0.015 <0.4 <0.4 <0.5<0.015 <0.015 0.11 0.11 <0.5Wrought Bal 22.369.00.723.32 0.80.030.0090.0050.08 0.30.18The L-PBF Inconel 625 samples were printed in two orientations i.e., Vertical (V), and Diagonal (D) as presented in Fig. 1. An island scanning strategy with a checkerboard pattern consisting of 5 × 5 mm squares was used. Each island alternated the scan direction perpendicular with-respect-to adjacent islands.

1. An island scanning strategy with a checkerboard pattern consisting of 5 × 5 mm squares was used. Each island alternated the scan direction perpendicular with-respect-to adjacent islands. Both vertical and diagonal samples were separated from the build plate via wire EDM and then subjected to various heat treatments. The V and D Inconel 625 samples were subjected to three heat-treatment temperatures of 700 ºC, 900 ºC, and 1050 ℃ for 1 hour. The rationale behind selecting these temperatures was to form the 𝛿 phase in the additively manufactured IN625; as shown in a TTT diagram provided by National Institute of Standards and Technology (NIST) [12]. Wrought Inconel 625 samples were subjected to the same heat-treatment temperatures for 40 hours. Some L-PBF specimens were not heat treated, i.e., left in the as-built condition. Heating ramps used during heat treatment were ~5 ºC/min. All specimens were subjected to air cooling to room temperature while remaining in the furnace.

Heating ramps used during heat treatment were ~5 ºC/min. All specimens were subjected to air cooling to room temperature while remaining in the furnace. Uncertainty in heat treatment temperatures is ± 5 ºC. For each sample, 5 microhardness measurements were taken on surfaces indicated in Fig. 1 using a Phase II 900-391D microhardness tester to be able to calculate average and standard deviation in measurements. Error bars on the plots represent the standard1028deviation. The indentations were made using a 1 kgf (9.8 N) load applied with a 15 secs dwell time.Figure 1: Vertically and diagonally built L-PBF Inconel 625 samples on build platform. Red arrows indicate the face used to measure microhardness.The samples were wrapped in aluminum foil and placed inside a 2.54 cm diameter vessel. The vessel was then placed into the reflector of the 10 MW reactor for 310 hours.

The vessel was then placed into the reflector of the 10 MW reactor for 310 hours. All samples faced a total flux of 6.61 x 1013 neutrons/cm²/s which corresponds to a fluence of 7.37 x 1019 neutrons/cm². The damage experienced by all samples was calculated to be 0.012 dpa. The microhardness measurements on radioactive samples were performed inside a hot cell equipped with manipulator arms as shown in Fig. 2.Figure 2: Hot cell with manipulator arms for performing microhardness measurements of radioactive Inconel 625 samples.1029Results and DiscussionsThe Vickers microhardness values of the as-built, pre-irradiated L-PBF Inconel 625 samples (built in vertical and diagonal directions) are shown in Fig. 3. For the vertical and diagonal samples, the microhardness values were measured to be 364 ± 1.175 HV and 359.25 ± 2.625 HV, respectively.

3. For the vertical and diagonal samples, the microhardness values were measured to be 364 ± 1.175 HV and 359.25 ± 2.625 HV, respectively. The two different build orientations showed minimal variation in microhardness indicating directional independency in hardness of the L-PBF Inconel 625. This observation is consistent with past research conducted by Wang et al. on Inconel 718 [13]. As-built L-PBF Inconel 625 samples showed superior microhardness compared to as-received wrought Inconel 625 which had a microhardness of 243.85 ± 1.15 HV. The higher microhardness in L-PBF samples is due to them having fine dendritic microstructure and high dislocation densities arising from high thermal gradients and cooling rates (~105 K/s) imposed by the fast-moving, high-heat-flux laser during the L-PBF process. The high cooling and solidification rates help in trapping the heavy atoms like Molybdenum (Mo) and Niobium (Nb) along interdendritic boundaries.

The high cooling and solidification rates help in trapping the heavy atoms like Molybdenum (Mo) and Niobium (Nb) along interdendritic boundaries. These heavy atoms located along sub domain boundaries restrict the sliding of dislocations under external stress which ultimately increases the hardness of the material. The lower hardness of the as-received wrought Inconel 625 is due to it possessing a more homogenous microstructure with equiaxed grains due to prolonged heat treatment and lower dislocation density. In addition, there is less segregation of heavy atoms like Mo and Nb along the grain boundaries since the amount of Nb in the wrought samples was less than that of the L-PBF Inconel 625 powder.

In addition, there is less segregation of heavy atoms like Mo and Nb along the grain boundaries since the amount of Nb in the wrought samples was less than that of the L-PBF Inconel 625 powder. The higher standard deviation in microhardness for the L-PBF samples may be due to them possessing a less homogeneous and highly textured microstructure and the micro segregation of heavy atoms like Nb and Mo in the microstructure.400V_As built350D_As builtW_As received)V H300(s s e n d r a h o r c i250200150M100500Figure 3: Microhardness of pre-irradiated Inconel 625 samples, including as-built/L-PBF vertical (V), as-built/L-PBF diagonal (D), and wrought (W). Error bars represent standard deviation.Performing heat treatments on the as-built L-PBF Inconel 625 samples was found to decrease microhardness. As presented in Fig. 4, increasing the heat-treatment temperature resulted in reduction of sample microhardness.

As presented in Fig. 4, increasing the heat-treatment temperature resulted in reduction of sample microhardness. Performing the same heat treatment process on wrought Inconel 625 showed different trends compared to L-PBF Inconel 625. The slight decrease in microhardness from the as-built state to 700 ℃ for the vertically-built L-PBF Inconel 625 sample is potentially due to a reduction in dislocation density and relaxation of micro strains. Interestingly,1030heat treating the L-PBF samples at 1050 ℃ brought their hardness values closer to that of the as- received wrought Inconel 625. The reduction in microhardness after heat treatment is most likely due to a decline in dislocation densities, solid state transformations, grain recrystallization or grain coarsening, and micro stress relaxation. Similar results were conveyed in past studies [5]. The microhardness increased when wrought Inconel 625 was heat treated at 900 ℃ and 1050 ℃.

Similar results were conveyed in past studies [5]. The microhardness increased when wrought Inconel 625 was heat treated at 900 ℃ and 1050 ℃. The likely reason for this is the onset of γ’’ and δ intermetallic phases forming in small amounts along sub-grain domain boundaries.400350)V H300(s s e n d r a h o r c i250200150M100500V _ A sb u i l tD _ A sb u i l tW _ A s b u i l tV _ 7 0 0o CD _ 7 0 0o CW _ 7 0 0 o CV _ 9 0 0o CD _ 9 0 0o CW _ 9 0 0 o CV _ 1 0 5 0 o CD _ 1 0 5 0 o CW _ 1 0 5 0o CFigure 4: Pre-irradiation microhardness measurements of as-built and heat-treated vertical (V) (dark grey) and diagonal (D) (light grey) L-PBF Inconel 625 samples and wrought (W) (dark red) Inconel 625 samples.The microhardness values of L-PBF and wrought Inconel 625 samples after thermal neutron irradiation are presented in Fig. 5.

5. It can be observed that neutron irradiation induced hardening occurred in all L-PBF and wrought Inconel 625 samples. Irradiation induced hardening was more prominent in the as-built diagonal L-PBF samples compared to the vertical-built samples. As-built diagonal L-PBF sample microhardness increased ~8% compared to only 1.2% in the as-built vertical L-PBF samples as presented in Fig. 6. It is hypothesized that the diagonal specimen layer-to-layer porosity and grain orientation are slightly different than that of the vertical specimens due to how they were oriented and manufactured (i.e., layer-wise scan strategy) during L-PBF. The smaller error bars shown in Fig. 5 for wrought Inconel 625, as compared to the L-PBF samples, suggest more homogeneity in microstructure and mechanical properties post irradiation in wrought Inconel 625. Heat treated vertical L-PBF and wrought samples at 1050 ℃ underwent slight softening as shown by negative percent change in Fig. 6.

Heat treated vertical L-PBF and wrought samples at 1050 ℃ underwent slight softening as shown by negative percent change in Fig. 6. This may be due to the dissolution of precipitates like γ’’ and δ, which primarily consist of Nb and Mo elements, that formed along sub domain grain boundaries. These precipitates restrict the movement of dislocations under external stress. Similar observations have been reported in past studies [14]–[16]. From Fig.

6, it is observed that all L-PBF Inconel 625 samples possessed better resistance to radiation induced hardening compared to all wrought Inconel 625 samples.1031a)b)Figure 5: Pre and post thermal neutron irradiation microhardness measurements of (a) all L-PBF Inconel 625 specimens built in vertical and diagonal orientations and (b) all wrought Inconel 625 samples.e g n a h Ct n e c r e P s s e n d r a h o r c i15 13 11 9 7 5 3 1 -1 -3 -5MV _ A sb u i l tD _ A sb u i l tV _ 7 0 0o CD _ 7 0 0o CV _ 9 0 0o CD _ 9 0 0o CV _ 1 0 5 0 o CD _ 1 0 5 0 o CW _ A sr e c e i v e dW _ 7 0 0 o CW _ 9 0 0 o CW _ 1 0 5 0o CFigure 6: Percent change in microhardness of as-built and heat-treated vertically (V) and diagonally (D) printed L-PBF and wrought (W) Inconel 625 samples after neutron irradiation.ConclusionsL-PBF Inconel 625 samples built in two different directions in the as-built and heat-treated conditions were successfully irradiated within a 10 MW nuclear reactor to study the phenomenon of neutron irradiation hardening.

The method of measuring microhardness of radioactive samples was established by using a hot cell and manipulator arms. Primary findings noted from this study are summarized below:The as-built vertical and diagonal L-PBF Inconel 625 samples had higher pre-irradiation hardness values compared to the as-received wrought Inconel 625 samples.1032The employed heat treatments used on the vertically- and diagonally-built L-PBF Inconel 625 samples decreased their microhardness as the temperature increased from 700 ℃ to 1050 ℃. The heat treatment process brought down the microhardness of L-PBF Inconel 625 imitating that of wrought Inconel 625.Even after modest neutron irradiation, all samples, except the vertically-built 1050 ℃ heat treated L-PBF and wrought 1050 ℃ heat treated samples, experienced neutron irradiation hardening.The as-built vertical L-PBF sample underwent less hardening (1.2%) than the as-built diagonal L-PBF sample (8%) indicating potential orientation effects.

As-received wrought samples faced 5.25% radiation induced hardening.The L-PBF Inconel 625 samples displayed more resistance to radiation hardening (lower percent change) as compared to wrought Inconel 625 samples (higher percent change).This study has provided more data indicating the structural integrity of components made from AM when exposed to full-spectrum nuclear radiation. Results provide insight into how one may minimize radiation hardening in additive manufactured materials for maintaining material property constraints. The results herein should assist engineers in selecting an appropriate heat treatment for additive manufactured nickel-based superalloys for increased radiation resistance. Results should increase confidence levels for adopting AM for building nuclear reactor components which perform the same or better than conventionally manufactured components.AcknowledgementThis material is based upon work supported by the U.S.

Department of Energy’s Office ofNuclear Energy under Award Number DE-NE0008865.DisclaimerThis report was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights. Reference herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof. The views and opinions of authors expressed herein do not necessarily state or reflect those of the United States Government or any agency thereof.References[1]A. F.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffects of Build Orientation and Heat Treatment on the Porosity Distribution and Morphology within Inconel 625 Fabricated via Laser Powder Bed FusionMohanish Andurkar1, Bart Prorok2, John Gahl3, Scott Thompson4,*1 Alan Levin Department of Mechanical and Nuclear Engineering, Kansas State University, Manhattan, KS 66506, United States of America.2 Department of Materials Engineering, Auburn University, Auburn, AL 36849, United States of America.

3Department of Electrical and Computer Engineering and University of Missouri Research Reactor (MURR), University of Missouri, Columbia, Missouri, 65211, United States of America.4Department of Mechanical and Aerospace Engineering and University of Missouri Research Reactor (MURR), University of Missouri, Columbia, MO 65211, United States of America.Corresponding author: smthompson@missouri.eduAbstractThe effects of build orientation, i.e., vertical, or diagonal (45º), and heat treatment on the porosity characteristics within Inconel 625 (IN625) fabricated via laser powder bed fusion (L- PBF) was experimentally investigated. Selected samples were heat treated at 1050 ℃ for 1-hour to promote evolution of pores. X-Ray Computed Tomography (XCT) was performed on samples to generate three-dimensional porosity maps. Volume Graphics (VG) software was used to inspect and quantify porosity distributions.

X-Ray Computed Tomography (XCT) was performed on samples to generate three-dimensional porosity maps. Volume Graphics (VG) software was used to inspect and quantify porosity distributions. Results indicate that build orientation and heat treatment influence measured porosity count. As-built (no heat treatment) sample microstructure was observed to have lower porosity count when compared to heat-treated samples. The vertically built sample was observed to have lower porosity relative to its diagonally built counterpart. The porosity morphology or diameter was observed to vary after heat treatment. On the other hand, the sphericity of pores was not affected by different build orientation and heat treatment.Keywords: laser powder bed fusion, X-ray CT, porosity measurements, pore morphologyIntroductionAdditive manufacturing (AM) continues to revolutionize how industry designs, manufactures, and distributes components to end users.

AM is versatile, flexible, highly customizable and, as such, has gained significant interest from academics and various industries. AM offers opportunities to build complex geometry parts with no additional tooling, reduction of process costs and potential to achieve zero waste manufacturing by maximizing material utilization, better control over the building process, on-demand manufacturing, and more [1]. Despite these capabilities of AM, its wide application for ‘printing’ high-value metal parts is currently restrained due to such parts having some level of microstructural defects such as porosity, voids, and residual stress. These defects can significantly degrade the structural integrity and performance of final components. Accurate detection, characterization, and prediction of these defects and irregularities have an important and immediate impact in the AM of fully dense parts [2]. ISO/ASTM 52900:2021 standard describes seven categories of AM.

ISO/ASTM 52900:2021 standard describes seven categories of AM. Powder bed fusion (PBF) is more accepted in industries due to several advantages it offers such as low cost, minimal support423structure requirements as the powder bed itself provides supports, larger material library, and the capability to recycle the unused powder [3].Porosity is defined as the ratio total pore volume to bulk material volume. Porosity can be considered as the Achilles heel of AM. Defect quantification and mitigation techniques have been an active research area in the metals AM field. Quantification of defects such as porosity in AM can be performed using x-ray computed tomography (XCT). XCT provides 3D information on internal defects by the reconstruction of a series of radiographs taken as the sample rotates over 180º or 360º [4].

XCT provides 3D information on internal defects by the reconstruction of a series of radiographs taken as the sample rotates over 180º or 360º [4]. The XCT technique is a widely applicable, powerful, and nondestructive inspection approach to evaluate and analyze geometrical and physical characteristics of materials, especially internal structures, and feature [5]. XCT has been used to characterize the size, shape, and volume of the pores in the sample. In the past, XCT analysis was used to correlate with process parameters used during the print to generate different porosity in parts [6]. A. Du Plessis provided a summary of pore types in AM metals [7]. Along with quantification, mitigation of these defects is equally critical to enhance the performance of the end user part. To minimize porosity in AM metals, it is important to carefully control the printing process.

To minimize porosity in AM metals, it is important to carefully control the printing process. This can be done by adjusting the laser power [8], [9], using high-quality metal powder [10], build orientation [11], and carefully selecting the printing parameters [12]. Post-processing techniques such as hot isostatic pressing (HIP) can also be used to reduce porosity and improve the mechanical properties of the final product. Although these studies were focused on varied materials, the porosity evolution in AM Inconel 625 (IN625) still needs to be comprehended in detail.The aim of this study is to investigate the effects of L-PBF build orientation and post- processing heat treatment on the porosity count, pore diameter, and sphericity in L-PBF IN625 samples. 3D porosity maps obtained via XCT technology will be provided.

3D porosity maps obtained via XCT technology will be provided. The goal is to provide a deeper understanding of the relationships between build orientation, heat treatment and porosity which should help in determining optimized process parameters for manufacturing highly dense components via L-PBF and subsequent heat treatments.Experimental MethodsAM IN625 cube samples were fabricated in a Concept Laser MLab 100R system. The printing was conducted in an argon atmosphere. An iterative method was used to determine optimized process parameters to print nearly full dense samples. The detailed optimize process parameters, scan strategy, powder composition used to print samples is explained in the study conducted previously [13]. Vertically and diagonally printed samples in the as-built and 1050 ℃- 1-hour heat treated condition were inspected. All samples were heat treated using a heating ramp of ~5℃/min and later air cooled in the furnace.

All samples were heat treated using a heating ramp of ~5℃/min and later air cooled in the furnace. The uncertainty in heat treatment temperatures is ± 5 ºC. The rationale behind using this heat treatment schedule was to investigate the effect of elevated temperature on the pore characteristics such as frequency, size, and sphericity.Porosity in a 3D XCT scan is identified as a low-density area (i.e., void space / air bubble) surrounded by higher density material. Raw XCT scan data for the part was reviewed and the region of interest to be analyzed was established and adjusted parameters which are used to define the pores. This parameter tuning phase increases the accuracy of both identification and quantification of each pore. A wealth of pore attribute data is calculated for each pore and includes pore count, size, and sphericity.Porosity maps and statistics were measured using a NSI X5000 225 XCT system available at Delphi Precision Imaging Corporation (Redmond, WA USA).

Sample part dimensions were approximately 5 mm x 5 mm x 10 mm in size. Samples were mounted on a rotating stage and424imaged using 255 kV and 160 𝜇A with 0.25 mm of copper (Cu) filter for beam hardening. A total of 5600 projections were taken using an oversampled sub-pixel scan technique with 5 integrated frames yielding a voxel size of 15 𝜇m. 3D XCT reconstruction was done with NSI EFX software; Analysis / Visualization was performed using commercial software VGSTUDIO MAX 2022 (Volume Graphic, Heidelberg, Germany). Porosity results can be visualized in the 3D CT scan model/2D slice views and individually color coded based on volume or other chosen attributes. This pore list can then be sorted and/or filtered based on a variety of attributes and can be exported as tabular data. The certainty of detected pores truly being gas pores (and not an artifact, surface void, inclusion, etc.) was set to be greater than or equal to 1 in the software.

The certainty of detected pores truly being gas pores (and not an artifact, surface void, inclusion, etc.) was set to be greater than or equal to 1 in the software. This ensured more accurate detection of pores with lighter gray contrast which indicates presence of gas. This study only focuses on the gas-entrapped pores and its quantitative analysis occurring in the microstructure of all samples.ResultsFigure 1 shows the total porosity count in all samples. The total porosity count in as-built vertical (V), diagonal (D), heat treated vertical, and diagonal samples are 186, 367, 1117, 1389, respectively. The as-built vertically printed sample has a lower porosity count compared to diagonally printed sample. This may be due to the presence of different heat transfer paths involving conduction between adjacent powder layers and Marangoni convention.

This may be due to the presence of different heat transfer paths involving conduction between adjacent powder layers and Marangoni convention. The difference in the thermal conductivity between layers oriented in vertical and diagonal directions also has influenced different porosity levels in the samples. 3D porosity maps of vertically and diagonally printed as-built and 1050 ℃- 1-hour heat treated IN625 samples are presented in Figures 2 and 3. As seen in Figs 2 and 3., there is a considerable influence of build orientation and heat treatment on the porosity counts in the AM IN625 samples.Fig 1. Porosity count in as-built vertical (V), as-built diagonal (D), 1050 ℃- 1h heat treated vertical printed, and 1050 ℃- 1h heat treated diagonal printed IN625 samples.425a) b)Fig 2. Porosity maps of a) as-built vertical and b) 1050 ℃- 1h heat treated vertical printed AM IN625 samples. Red points represented pores.a) b)Fig 3.

Porosity maps of a) as-built vertical and b) 1050 ℃- 1h heat treated vertical printed AM IN625 samples. Red points represented pores.a) b)Fig 3. Porosity maps of a) as-built diagonal and b) 1050 ℃- 1h heat treated diagonal printed AM IN625 samples. Red points represented pores.Pore diameter distribution found within as-built and 1050 ℃- 1-hour heat treated vertically printed samples is presented in Figure 4. Post 1050 ℃- 1-hour heat treatment, the pore diameters in the vertically printed samples were observed to increase. The average pore diameter for the as- built and 1050 ℃- 1-hour heat treated vertical samples are 60.226 ± 10 and 67.35 ± 10.2 𝜇m, respectively. The formation of pores and the increased pore diameter due to heat treatment426presumably resulted from the temperature-driven equilibrium diffusion and increase gas (argon) pressure inside of the pores. Similar observations were reported in previous studies [14], [15].a)b)Fig 4.

Similar observations were reported in previous studies [14], [15].a)b)Fig 4. Pore diameter distribution (in 𝜇m) in a) as-built vertical and b) 1050 ℃- 1-hour heat treated vertical printed AM IN625 samples. Average and standard deviation are represented with mu and sigma values, respectively.Pore diameter distribution found within as-built and 1050 ℃- 1-hour heat treated diagonally printed sample is presented in Figure 5. A similar trend compared to vertically printed samples was observed in the diagonally printed samples. The average pore diameter for the as-built and 1050 ℃- 1-hour heat treated diagonal samples are 59 ± 8.13 and 68.48 ± 10.59 𝜇m respectively. This indicates regardless of build orientation, heat treatment resulted in the expansion of the pores.

This indicates regardless of build orientation, heat treatment resulted in the expansion of the pores. In addition, it can be observed that pore diameter distribution in the as-built vertically and diagonally samples showed very slight variation in the average pore diameter value, indicating427minimal effect of the build orientation on the average pore diameter. This is indicated by the similarity in the mu values shown in Figures 4 and 5.a)b)Fig 5. Pore diameter distribution (in 𝜇m) in a) as-built diagonal and b) 1050 ℃- 1-hour heat treated diagonal printed AM IN625 samples. Average and standard deviation are represented with mu and sigma values, respectively.Pore sphericity within the as-built and 1050 ℃- 1-hour heat treated vertically printed samples is presented in Figure 6. As presented in Fig. 6., the effect of heat treatment on the sphericity of the pores is not noticeable at all.

As presented in Fig. 6., the effect of heat treatment on the sphericity of the pores is not noticeable at all. The average pore sphericity in the as-built and 1050 ℃- 1-hour heat treated vertically built samples are 0.94 ± 0.2. The mu and sigma values in Figs. 6 and 7 represent average and standard deviation values. The red curve represents the distribution curve of the data.428a)b)Fig 6. Pore sphericity distribution in a) as-built vertical and b) 1050 ℃- 1-hour heat treated vertical printed AM IN625 samples. Average and standard deviation are represented with mu and sigma values, respectively.Pore sphericity within the as-built and 1050 ℃- 1-hour heat treated diagonally printed samples is presented in Figure 7. Similar observation was made in diagonally built samples compared to the vertically built samples. The average pore sphericity in the as-built and 1050 ℃- 1-hour heat treated diagonally built samples are 0.94 ± 0.2. In addition, from Figs.

The average pore sphericity in the as-built and 1050 ℃- 1-hour heat treated diagonally built samples are 0.94 ± 0.2. In addition, from Figs. 6 and 7, it can be observed that the difference in the build orientation has a negligible effect on the average pore sphericity value.429a)b)Fig 7. Pore sphericity distribution in a) as-built diagonal and b) 1050 ℃- 1-hour heat treated diagonal printed AM IN625 samples. Average and standard deviation are represented with mu and sigma values, respectively.Conclusions The present study investigated the effects of the build orientation and post processing heat treatment on the porosity characteristics in IN625 samples fabricated using laser powder bed fusion (L-PBF) additive manufacturing (AM). Samples were printed at vertical and diagonal orientation and used in as-built, and heat treated at 1050 ℃ for 1-hour conditions. Major conclusions are summarized as follows:1.

Samples were printed at vertical and diagonal orientation and used in as-built, and heat treated at 1050 ℃ for 1-hour conditions. Major conclusions are summarized as follows:1. Build orientation and heat treatment had an influence on porosity count. The as-built vertical sample showed a lower porosity count compared to the as-built diagonal sample.430The as-built vertical sample had 186 pores compared to 367 pores in the as-built diagonal sample.2. Heat treatment at 1050 ℃-1h, increased porosity count in both orientation-built samples. The 1050 ℃-1h heat treated vertical and diagonal samples had 1117 and 1389 porosity counts respectively.3. Pores size increased in vertical and diagonal printed samples post 1050 ℃-1h heat treatment indicating the pressure increase in the pore.4. However, neither build orientation nor heat treatment had a noticeable effect on the sphericity of pores in all samples. The average sphericity in all samples was observed to be 0.94.

However, neither build orientation nor heat treatment had a noticeable effect on the sphericity of pores in all samples. The average sphericity in all samples was observed to be 0.94. Acknowledgements4. However, neither build orientation nor heat treatment had a noticeable effect on the sphericity of pores in all samples. The average sphericity in all samples was observed to be 0.94. AcknowledgementsThis material is based upon work supported by the U.S. Department of Energy’s Office ofReferences[1]S. A. M. Tofail, E. P. Koumoulos, A. Bandyopadhyay, S. Bose, L. O’Donoghue, and C. Charitidis, “Additive manufacturing: scientific and technological challenges, market uptake and opportunities,” Materials Today, vol. 21, no. 1, pp. 22–37, 2018, doi: https://doi.org/10.1016/j.mattod.2017.07.001.[2]A. Mostafaei et al., “Defects and anomalies in powder bed fusion metal additive manufacturing,” Current Opinion in Solid State and Materials Science, vol. 26, no. 2, p.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEffects of build parameters on compression properties for ULTEM 9085 parts by fused deposition modelingKrishna P. Motaparti 1, Gregory Taylor 1, Ming C. Leu 1, K. Chandrashekhara 1, James Castle 2 and Mike Matlack 2 1 Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, Missouri, USA 2 The Boeing Company, St Louis, Missouri, USAAbstractIt has been observed by various researchers that parts fabricated by the Fused Deposition Modeling (FDM) process have anisotropic properties. The research presented in the present paper was aimed to study the compression properties of FDM parts and to comprehend their dependence on build parameters.

The research presented in the present paper was aimed to study the compression properties of FDM parts and to comprehend their dependence on build parameters. In this study Ultem 9085 was used as the material to fabricate both solid and sparse-build coupons with variations in build direction, raster angle and air gap. A full factorial experimental design was used to study the individual and combined effects of these build parameters on the mechanical properties of the coupons. The mechanical properties studied include compressive yield strength, compressive modulus, compressive strength/mass ratio, and compressive modulus/mass ratio. Besides the obtained test data, qualitative observation and reasoning was used to help understand how the compression properties are affected by the build parameters.1. IntroductionAdditive manufacturing (or 3D printing) is a method of fabricating three-dimensional prototypes or parts through the deposition of layers upon layers of material.

IntroductionAdditive manufacturing (or 3D printing) is a method of fabricating three-dimensional prototypes or parts through the deposition of layers upon layers of material. Currently, additive manufacturing has spread into numerous industries and is quickly becoming a widely accepted manufacturing technique for more than just prototypes. Contrary to subtractive manufacturing techniques, additive manufacturing has a greatly reduced material requirement due to the layer- by-layer building process. Additive manufacturing has a variety of building techniques, but one in particular, Fused Deposition Modeling (FDM), is often utilized for the fabrication of thermoplastic components. The fabrication is controlled by machine code generated from a CAD model in STL (STereo Lithography) format. With the build parameters and STL file, an FDM machine fabricates the physical part through an extrusion of the thermoplastic material from a nozzle head.

With the build parameters and STL file, an FDM machine fabricates the physical part through an extrusion of the thermoplastic material from a nozzle head. In the FDM machine of this study, the build platform has motion in the vertical Z direction whereas the nozzle head has motion constrained to the horizontal plane (i.e. in X and Y directions).Currently, there is a need to understand the mechanical behavior of FDM specimens and researchers have been studying the dependence on build parameters such as build direction, raster orientation, raster width, layer thickness, oven temperature, etc. [1-10]. Schopper et al. [1] studied the effect of build direction on the compression properties of FDM specimens and they reported that the yield strength and compressive modulus of specimens built in the horizontal direction were higher in comparison to the specimens built in the vertical direction. Also, according to Bagsik et al.

Also, according to Bagsik et al. [2], the tensile properties of specimens made of Ultem 9085 were1964higher in all build directions using a negative air gap. Lee et al. [3] conducted a case study by using Taguchi method to investigate the build parameters in order to achieve optimum performance for a compliant ABS prototype. Lee et al. [4] compared the compressive strengths of parts made by FDM, inkjet printing, and nano composite deposition system (NCDS). They reported that the parts made by FDM had high compressive strength in comparison to the other processes. Ognzan et al. [5] studied the effects of layer thickness, deposition angle, and infill percentage on the maximum flexural force in FDM specimens made of polylactic acid (PLA). They concluded that layer thickness has the maximum effect on the flexural strength followed by the interaction between deposition angle and infill percentage.

They concluded that layer thickness has the maximum effect on the flexural strength followed by the interaction between deposition angle and infill percentage. A relationship between the total costs of FDM parts and their mechanical properties was established by Rauta et al. [6] to enable the engineers to decide on proper build orientations so that FDM parts can be fabricated with good mechanical properties at low manufacturing costs. The effects of raster angle, oven temperature, and raster width on the properties of FDM parts made of ABS were studied using a bacterial foraging technique by Panda et al. [7] in order to determine optimal build parameter settings to achieve good strength. Rayegani et al. [8] used the group method for prediction purposes and developed a functional relationship between build parameters and the part’s tensile strength for the FDM process. During the production of components using FDM, factors like build time and surface roughness play an important role.

During the production of components using FDM, factors like build time and surface roughness play an important role. Several studies were conducted to optimize these factors using different build parameters. Anitha et al. [9] used Taguchi method to study the various build parameters used in FDM that affect the quality of the fabricated part. The quality was measured in terms of the surface roughness of the part and it was found that layer thickness, raster width, and speed of deposition influence the part quality with layer thickness having the maximum effect. Similar experiments conducted by Vasudevarao et al. [10] revealed that part orientation also affects surface finish but air gap does not have significant influence on surface quality.

Similar experiments conducted by Vasudevarao et al. [10] revealed that part orientation also affects surface finish but air gap does not have significant influence on surface quality. The technical literature mostly contains studies that analyze the measured data from experiments, and there have been very few qualitative reasoning studies for the observed part properties.The main objective of this research is to study how various build parameters including build direction, raster angle, and air gap affect the compressive modulus and yield strength of FDM produced Ultem 9085 specimens with solid and sparse build. The study includes a full- factorial design of experiments to help understand how much of an effect, if any, the build parameters have on the compression properties.2. Material and Methods2.1. Additive manufacturing machine and materialIn this study, all specimens were fabricated with a Fortus 400mc (Stratasys).

Material and Methods2.1. Additive manufacturing machine and materialIn this study, all specimens were fabricated with a Fortus 400mc (Stratasys). The machine is capable of building components within a 406 mm x 356 mm x 4.6 mm (16” x 14” x 16”) build envelope and with a ± 0.0015 mm/mm accuracy. While the Fortus 400mc is able to build with many thermoplastic materials such as ABS, Ultem, Polycarbonate (PC), and PPSF, this study used Ultem 9085 as the specimen material. Ultem 9085 is regarded as a high- performance thermoplastic with high temperature capacity and good strength-to-weight ratio. With the fabrication of Ultem 9085 specimens, a breakaway support material was used. Fabrication of the specimens consisted of the following steps:29651. Pre-processing: Three-dimensional models of the test specimens are modeled in CAD software (SolidWorks) and exported as a Stereo Lithography (STL) file.

Pre-processing: Three-dimensional models of the test specimens are modeled in CAD software (SolidWorks) and exported as a Stereo Lithography (STL) file. The STL file is taken as input into Statasys Insight 9.1 software where the fabrication tool path is generated from user-specified build parameters. Figure 1 shows a potential schematic for a rectangular layer of FDM component.2. Fabrication: After pre-processing, the STL file is sent to the Fortus 400mc using Stratasys Insight and Control Center Job Processing and Management software. The specimen is then fabricated using the FDM process.3. Post-processing: After fabrication, the support structure is mechanically removed from the model and the specimen is finished.Figure 1. FDM layer schematic in Insight software showing different build parameters2.2. Specimen fabricationFor this study, the Ultem 9085 specimens were built with two build directions, horizontal and vertical.

Specimen fabricationFor this study, the Ultem 9085 specimens were built with two build directions, horizontal and vertical. For the horizontal-build, specimens are fabricated with the build direction parallel to the compression load to be applied later. For the vertical-build, specimens are fabricated with the build direction perpendicular to the compression load to be applied. Raster angle varied with either (45°, -45°) and (0°, 90°). Air gap varied with -0.00635 mm (-0.00025”), -0.0127 mm (- 0.0005”), or -0.01905 mm (-0.00075”) for solid specimens but was held at a constant 2.54 mm (0.1”) for the sparse specimens. Raster width and contour width were both held at a constant 0.508 mm (0.02”) for all specimens. Specimens were fabricated with a 38.1 mm x 38.1 mm (1.5” x 1.5”) cross-section and 25.4 mm (1”) height as shown in Figure 2. Compression tests were performed with a 1.27 mm/min (0.05 in/min) loading rate and were stopped when the specimen had over 10% strain.

Compression tests were performed with a 1.27 mm/min (0.05 in/min) loading rate and were stopped when the specimen had over 10% strain. Effective modulus was calculated as stress divided by strain in the elastic region and effective yield strength was calculated as strength at 0.2% offset.For all specimens, compression tests were conducted on an Instron 5980 which records the load and extension for each test specimen. Instron software was then used to calculate and plot the stress vs. strain curve for each specimen. Stress is defined as the compressive load per unit area of the specimen cross-section. Compressive strain is defined as the ratio of compressive deformation to the gage length of the test specimen.3966Figure 2. Dimensions, horizontal-build specimen, and vertical-build specimen used in this study2.3.

Dimensions, horizontal-build specimen, and vertical-build specimen used in this study2.3. Design of experiments – factors and levelsIn design of experiments, independent variables, also known as factors, and variations for these variables, also known as levels, are established. This study splits the specimens into two experiments, solid and sparse specimens.For the solid specimen case, three factors including build direction, raster angle, and air gap were used (Table 1). Build direction has two levels, horizontal and vertical. Raster angle has two levels, (0°, 90°) and (45°, -45°). Air gap has three levels, -0.00635 mm (-0.00025”), -0.0127 mm (-0.0005”) and -0.01905 mm (-0.00075”). While increasing the magnitude of the negative air gap in the solid specimens will generally increase the strength, the negative air gap will greatly decrease the surface quality of the specimens when the air gap reaches a certain limit (- 0.01905 mm for the Fortus 400mc).

A full-factorial experiment was performed using the 12 combinations with replications of 5 samples for each combination.Table 1. Factors and levels for solid build styleFactorsBuild directionLevels (5 Replications) -1 1 0 Vertical - HorizontalRaster angle (degrees/degrees)0/9045/-45Air gap (mm)0.006350.01270.01905In the sparse case, two factors including build direction and raster angle were used (Table 2). Unlike the solid case, the air gap was fixed at 2.54 mm (0.1”). Build direction has two levels, horizontal and vertical. Raster angle has two levels, (0°, 90°) and (45°, -45°). A full-factorial experiment was performed using the 4 combinations with replications of 5 samples for each combination.Table 2. Factors and levels for sparse build styleFactorsBuild directionLevels (5 Replications) -1 1 0 Vertical - HorizontalRaster angle (degrees/degrees)0/9045/-4549673.

Factors and levels for sparse build styleFactorsBuild directionLevels (5 Replications) -1 1 0 Vertical - HorizontalRaster angle (degrees/degrees)0/9045/-4549673. Experimental ResultsTwo types of experiments were performed in this study. The first experiment was performed to study the effects of build direction and temperature on both solid and sparse compression properties of FDM specimens. The second experiment was aimed at understanding the effects build parameters including build direction, raster angle, and air gap on the compression properties of FDM specimens. The second experiment was further divided into solid and sparse specimens.3.1. Effects of build style and temperatureFive specimens were tested for each combination of solid/sparse, build direction, and temperature. Build styles included horizontal-build and vertical-build while temperatures included 24°C (75°F), 82°C (180°F), and 121°C (250°F).

Build styles included horizontal-build and vertical-build while temperatures included 24°C (75°F), 82°C (180°F), and 121°C (250°F). The obtained compressive modulus, yield strength, specific modulus, and specific yield strength of the specimens were averaged and plotted along with standard deviation.From compression test results shown in Figure 3, at all three temperatures [24°C (75°F), 82°C (180°F), 121°C (250°F)], the yield strength, compression modulus, strength/mass ratio, and modulus/mass ratio of the horizontal-build specimens (S-H and SP-H) are higher in comparison to the corresponding vertical-build specimens (S-V and SP-V). The specimens with S-H build style (solid specimen with horizontal-build) exhibits the highest strength and modulus at each respective temperature. These results clearly indicate that both solid/sparse build and build direction affect the compression properties of FDM specimens.Figure 3.

These results clearly indicate that both solid/sparse build and build direction affect the compression properties of FDM specimens.Figure 3. Comparison of compression properties for solid and sparse specimens with different build directions for three temperatures (S indicates a solid specimen, SP indicates a sparse specimen, H indicates horizontal-build direction, and V indicates vertical-build direction)5968The results show the yield strength of the solid specimen decreases with increase in temperature. In the case of horizontal-build specimens, the yield strength decreases more than the vertical-build specimens for each respective temperature. This indicates that the vertical- build specimens display a relatively higher resistance to temperature in comparison to the horizontal-build specimens, whereas horizontal-build specimens display relatively higher strength.Figure 4. Decrease in yield strength of horizontal-build and vertical-build solid specimens3.2.

Decrease in yield strength of horizontal-build and vertical-build solid specimens3.2. Effects of build direction, raster angle and air gapA second type of experiment was designed to conduct compression tests on both solid and sparse specimens with varying build parameters including build direction, raster angle, and air gap (air gap constant for sparse specimens). A full-factorial design of experiment was used for both solid and sparse specimens and the effects of the parameters were studied.3.2.1. Solid-build styleCompression tests were conducted on the FDM fabricated solid specimens. A total of 60 specimens were tested for 12 different combinations of solid build parameters and 5 replications for each set of parameters. The study investigated compression properties including modulus, yield strength, specific modulus, and specific yield strength. The full-factorial experiment was used for the determination of the main effects and the interactions among the parameters.

The full-factorial experiment was used for the determination of the main effects and the interactions among the parameters. Main effect is defined as the effect of an independent variable on a response variable averaging across different levels of other independent variables. Independent variables include build direction, raster angle, and air gap for the solid specimen case. If the effect of one independent variable on the response variable is dependent on the value of another independent variable, those two variables are said to exhibit interaction. Yield strength was the only response variable considered during the experiment. All results were analyzed with JMP 11. Results from the statistical software are given in Table 3. The p-value [19] defines the level of significance within a statistical test and represents the probability of a factor affecting the outcome (response variable).6969Table 3.

The p-value [19] defines the level of significance within a statistical test and represents the probability of a factor affecting the outcome (response variable).6969Table 3. Effects table from JMP 11 for solid specimen caseFactorsSum of SquaresF Ratiop-valueBuild direction 4775.5037 8392.691 <.0001 343.6323 603.9153 <.0001 Raster angle 0.0816 3.4946 Air gap <.0001 438.708 Build*Raster 0.1019 2.7547 Build*Air gap 0.2699 1.2385 Raster*Air gap6.9546 249.6281 1.5674 0.7047For this statistical experiment, if the p-value is less than 0.05 (for a significance level of 95%), the corresponding factor has a significant effect on the response variable. All p-values less than 0.05 are highlighted in red to indicate main effects and/or interaction. The effects table indicates an interaction between build direction and raster angle, which means that the effect of raster angle on the response variable is dependent on the build direction used for the specimen.

The interaction of these parameters is described below in detail.Interaction of build direction and raster angle: The two different raster angles evaluated in this experiment included (0°, 90°) and (45°, -45°) for both horizontal-build and vertical-build solid specimens. The results are shown in Figures 6 and 7.Figure 6. Interaction of build direction and raster angle on yield strength of solid specimensFigure 7. Sample compression stress-strain curves for horizontal-build and vertical-build solid specimens (air gap = -0.00635mm)7970Horizontal build: According to the graphs in Figures 6 and 7, the horizontal-build specimens with (0°, 90°) and (45°, -45°) raster angles have very small differences in the yield strength, and the trend remains the same for different values of air gap.

This occurs because in the case of horizontal-build specimens, the applied load is perpendicular to the layer in which the rasters are present, and the internal structure for resisting the external load is essentially the same for different raster angles; refer to Figure 8.Figure 8. Physical models for horizontal-build compression solid specimens with (0°, 90°) and (45°, -45°) raster anglesVertical build: From the graphs in Figures 6 and 7, the vertical-build specimens with (0°, 90°) raster angles exhibit about 12% higher yield strength in comparison to the specimens with (45°, -45°) raster angles. The trend remains the same for different air gaps. This indicates that the specimens built with (0°, 90°) raster angle offer more resistance to deformation in comparison to the specimens built with (45°, -45°) raster angle. Figure 9 illustrates the (0°, 90°) vs. (45°, -45°) raster angle in vertical-build solid specimens.Figure 9.

Figure 9 illustrates the (0°, 90°) vs. (45°, -45°) raster angle in vertical-build solid specimens.Figure 9. Physical models for vertical-build solid specimens with (0°, 90°) and (45°, -45°) raster anglesEffect of air gap: The three air gaps evaluated in this experiment included -0.00635 mm (-0.00025”), -0.0127 mm (-0.0005”) and -0.01905 mm (-0.00075”) for both horizontal-build and vertical-build solid specimens. The results are shown in Figures 10 and 11.8971Figure 10. Effect of air gap on yield strength for solid specimensFigure 11. Stress-strain curves at different air gaps for horizontal-build and vertical-build specimens, where the raster angle was (0°, 90°)Horizontal build: From Figures 10 and 11 and the effects table in Table 3, it can be seen that in the case of horizontal-build solid specimens, the effect of air gap is not significant with a p-value > 0.05. The trend remains the same for both (0°, 90°) and (45°, -45°) raster angles.

The trend remains the same for both (0°, 90°) and (45°, -45°) raster angles. This indicates that the variation in air gap between the rasters does not have a significant effect on the failure due to inter-layer sliding in solid specimens.Vertical build: The effect of air gap in vertical-build is also not significant and the trend remains the same for (0°, 90°) and (45°, -45°) raster angles. As indicated by the effects table, the influence of air gap on the yield strength is not significant (p-value > 0.05), i.e., there is no main effect of air gap for vertical-build. This indicates that the variation in air gap between the rasters also does not have a significant effect on the failure due to buckling in solid specimens.99723.2.2. Sparse-build styleCompression tests were conducted on the FDM fabricated sparse specimens. A total of 20 specimens were tested for 4 different combinations of sparse-build parameters and 5 replications for each set of parameters.

A total of 20 specimens were tested for 4 different combinations of sparse-build parameters and 5 replications for each set of parameters. The air gap was held at a constant 2.54 mm (0.1”) for all sparse samples. A full-factorial experiment was conducted to determine the statistical significance of the build direction and raster angle. The effects table created in JMP 11 shows the main effects and interaction of the two parameters (Table 4).Table 4. Effects table from JMP 11 for sparse specimen caseFactorBuild direction Raster angle Build*RasterSum of Squares 1062.88 138.33 175.23F Ratio814.62 106.02 134.3p- Value<.0001 <.0001 <.0001From the P-values in Table 4, the build direction and raster angle and their interaction each have a significant effect (p-value < 0.05) on the yield strength.

The interaction of these parameters is described below in detail.Interaction of build direction and raster angle: From Figures 12 and 13, the strength of specimens built in horizontal direction is 19% to 40% higher than that in vertical direction for sparse specimens. This behavior is similar to the one observed in the case of solid specimens, and it is expected since the layers in vertical-build specimens fail due to buckling of layers, whereas in the case of horizontal-build specimens, the failure occurs due to inter-layer sliding. Similar to solid-build style, a change in raster angle on horizontal-build sparse specimens shows only a small difference in the yield strength of the sparse specimen. This is expected as the applied compression force acts perpendicular to the layer in which the rasters are present. Similar to solid specimens, the internal structure for resisting the external load is essentially the same for different raster angles.

Similar to solid specimens, the internal structure for resisting the external load is essentially the same for different raster angles. However, in the case of vertical-build, the effect of raster angle is more pronounced as seen in Figure 12. The specimens with (0°, 90°) raster angle have about 30% higher yield strength in comparison to the specimens with (45°, -45°) raster angle. This indicates that the (0°, 90°) rasters produce relatively stiffer specimens compared to (45°, -45°) raster specimens.1097350) a P M( h t g n e r t s d e Yli40302010Raster angle0-9045-450HorizontalVerticalBuild directionFigure 12. Effect of raster angle for horizontal-build and vertical-build sparse specimensFigure 13. Sample stress-strain curves for horizontal-build and vertical-build sparse specimens3.2.3.

Effect of raster angle for horizontal-build and vertical-build sparse specimensFigure 13. Sample stress-strain curves for horizontal-build and vertical-build sparse specimens3.2.3. Failure of vertical-build sparse specimensFrom the above results, the mechanical properties of vertical-build sparse specimens with (0°, 90°) raster angle are higher in comparison to the specimens with (45°, -45°) raster angle in compression tests. In the case of (0°, 90°) raster angle, only the rasters deposited perpendicular to the loading surface will resist the load, but in the case of (45°, -45°) raster angle, the applied load is taken by all the rasters. This behavior can be explained by examining the specimens during the compressive testing as shown in Figure 14. Figure 15 represents the load vs. deformation relationship for (0°, 90°) and (45°, -45°) raster angles.

Figure 15 represents the load vs. deformation relationship for (0°, 90°) and (45°, -45°) raster angles. According to the graph, (0°, 90°) raster specimens exhibit about 40% higher yield strength and stiffness in comparison to (45°, -45°) specimens. In the case of (0°, 90°) raster specimens, the deposited vertical rasters act as struts and resist the deformation until a load of ~43 kN, and the specimen suddenly fails internally due to buckling at ~2 mm deformation. In the case of (45°, -45°) raster specimen, the failure does not occur suddenly. The sparse specimen created with (45°, -45°) raster angle undergoes more deformation before failure occurs. The structure reaches a deformation of ~2 mm at a load of ~26 kN, which is about 40% less than the failure load of (0°, 90°) raster specimen. All of the tested samples for each set of specimens exhibit the same behavior.

All of the tested samples for each set of specimens exhibit the same behavior. The11974failure in the (45°, -45°) raster specimen occurs on the contour of the specimen upon continuous loading as shown in Figure 15. Thus, the (0°, 90°) raster specimen is comparatively stiffer and stronger compared to the (45°, -45°) raster specimen.Figure 14. Gradual failure of vertical-build sparse specimens with (0º, 90º) and (45º, -45º) raster angles in compression testsFigure 15. Load vs. deformation for vertical-build sparse specimens in compression tests4. ConclusionThe FDM process was used to fabricate solid and sparse Ultem 9085 specimens. The specimens were varied in build parameters, and compressive properties including modulus and yield strength were studied. A full-factorial experiment was utilized to determine the effects of the build parameters including build direction, raster angle, and air gap.

A full-factorial experiment was utilized to determine the effects of the build parameters including build direction, raster angle, and air gap. From the statistical data,12975it can be concluded that for both solid and sparse specimens, the interaction between build direction and raster angle is the most significant factor in the resulting differences in yield strength between specimens.The compression test results indicate that the horizontal-build direction exhibits 15-40% higher compressive strength in comparison to vertical-build direction for both solid and sparse build specimens. The compressive yield strength of vertical-build solid and sparse specimens is 12-30% higher with (0°, 90°) raster angle in comparison to (45°, -45°) raster angle; however, the difference in compressive yield strength due to different raster angles was very small for horizontal-build specimens.

The effect of negative air gap on the compressive yield strength was determined to be not statistically significant (p-value > 0.05) for solid specimens.5. AcknowledgementsThis research is sponsored by the Industrial Consortium of the Center for Aerospace Manufacturing Technologies (CAMT) at Missouri University of Science and Technology. Special thanks are due to Michael Hayes of Boeing and Timothy Schniepp and Chris Holshouser of Stratasys for their valuable contributions to the project.6. References1. Bagsik, A., Schoeppner, V., “Mechanical properties of fused deposition modeling parts manufactured with ultem 9085,” Proceedings of the ANTEC, Plastics: Annual Technical Conference Proceedings, ANTEC 2011.2. Bagsik, A., Schoeppner, V., Klemp, E., “FDM part quality manufactured with ultem 9085,” 14th International Scientific Conference on Polymeric Materials, Halle (Saale), 2010.3.

As opposed to traditional manufacturing methods, a part is manufactured layer by layer from 3D CAD models in AM. Though the vision of AM is impressive, there are many challenges that are hindering the widespread use of these complex parts. One of the challenges in these materials is defects grow and their orientation during the manufacturing process. In this paper we are going to investigate the effect of Build Parameters and their effects on the mechanical and electrical properties of the additively manufactured heterogeneous material system. We will study the electrical properties to find out the material state of additively manufactured part.Additive manufacturing (AM) revolutionized manyindustries,IntroductionAdditive manufacturing(AM), a formal term for Rapid Prototyping(RP) is used in a variety of industries to describe a process for rapidly creating a system or part representation before final release or commercialization.

In other words, the emphasis is on creating something quickly and that the output is a prototype or basis model from which further models and eventually the final product will be derived [1]. There are a variety of individual processes which differ in their method of layer manufacturing. Individual processes will differ depending on the material and machine technology used. Material extrusion is one of the most common and commercial method used for Additive Manufacturing. The printer that was used for this paper follows the principle of ‘Fused Deposition Modelling(FDM)’. This approach extrudes the material through a nozzle in a controlled manner. FDM uses a heating chamber to liquefy polymer that is fed into the system as a filament. The filament is pushed into the chamber by a tractor wheel arrangement and it is this pushing that generates the extrusion pressure as shown in Figure 1. The main drawback to using this technology is the build speed.

The main drawback to using this technology is the build speed. The inertia of the plotting heads means that the maximum speeds and accelerations that can be obtained are somewhat smaller than other systems. Furthermore, FDM requires material to be plotted in a point-wise, vector fashion that involves many changes in direction. [1]1536Figure 1 :- Fused Deposition Molding [2]For practical applications, these 3D-printed parts must tolerate significant amounts of mechanical, thermal, environmental etc. loads and hence it is of great importance to determine the required strengths, and at the very least, the physical properties of these 3D-printed parts should be comparable to those manufactured by conventional methods [3].

Due to the nature of the process that builds the parts layer by layer, the mechanical properties of the part depend on the material used and processing parameters which include the raster orientation, extrusion temperature, bed temperature, infill percentage, patterns and many others. For the following study, bed temperature and infill percentage are chosen as variables to determine their effect on the material performance. To characterize these materials, mechanical and dielectric tests (using Broadband dielectric spectroscopy (BbDS) are performed on the printed parts.BbDS is a well- established tool for dielectric material characterization which has been used in polymer industries for a long time, e.g., in composite manufacturing this method is used to monitor the curing process.

A material system is comprised of multiple polarization mechanisms such as ionic (molecular), dipolar (orientational), electronic, interfacial (Maxwell–Wagner– Sillars) polarization and hopping charge polarization. The following Figure 2 shows the different types of polarization and their effect on the dielectric response and its corresponding effective frequency range. BbDS is the interaction of electromagnetic waves with matter in the frequency range from a lower value of 10-6 Hz to a higher frequency of 1012 Hz.

BbDS is the interaction of electromagnetic waves with matter in the frequency range from a lower value of 10-6 Hz to a higher frequency of 1012 Hz. This dynamic range contains information about the molecular and collective dipolar fluctuation; charge transport and polarization effects occur at inner and outer boundaries in the form of different dielectric properties of the material under study [4].1537Figure 2:- Effect of different charge displacement mechanisms on dielectric response and their corresponding effective frequency rangePolar molecules are contained inside of heterogeneous materials and in the presence of an applied electric field it will polarize the material by orienting the dipole moments of polar molecules and charge accumulation at the interfaces of dissimilar materials. The polarization of a linear dielectric is given as:𝑃 = ℇ0 𝜒 𝐸here, 𝜒 = tensor of dielectric susceptibility ℇ0 = dielectric permittivity of vacuum i.e., 8.854*10e-12 F/m 𝐸 = Electric Field.

The polarization of a linear dielectric is given as:𝑃 = ℇ0 𝜒 𝐸here, 𝜒 = tensor of dielectric susceptibility ℇ0 = dielectric permittivity of vacuum i.e., 8.854*10e-12 F/m 𝐸 = Electric Field. Now, based on Maxwell’s laws of electromagnetism, we have the equation,(2) This is also known as Gauss Law, Where, 𝛻 ∙ = Divergence operator, 𝐷 = Electric∇ ∙ 𝐷 = 𝜌Displacement Field and 𝜌 = Electric charge volume densityThe relation between the dielectric displacement and electric field is given as 𝐷 = ℇ0𝐸 + 𝑃From (1),𝐷 = ℇ0𝐸 + ℇ0𝜒𝐸𝐷 = ℇ0ℇ𝑟𝐸In this case, ℇ𝑟 = (1 + 𝜒) is known as relative permittivity. Finally, we get 𝐷 = ℇ𝐸 where ℇ = ℇ0ℇ𝑟 = ℇ′ − ℇ" is defined as complex permittivity in which ℇ′ is the real part of the complex permittivity and ℇ" is the imaginary part, in other words the dielectric loss.

Thus, we can now plot two graphs, one for the real part of the permittivity and other for the imaginary part, both of which corresponds to the logarithmic values of the frequencies.The schematic of the setup is shown below in Figure 3. The arrangement is like a parallel plate capacitor, where in one end of the electrode an alternating signal is supplied and the other end of the electrode is grounded.1538(1)(3)(4)(5)Figure 3:- Schematic of the dielectric response setupAn AC voltage can be represented by equation 6.U(t) = U0 cos(ωt) = Re(U∗ exp(iωt))Where U∗ = U0, is the amplitude of the signal applied, ω is the frequency of the appliedsignal.

The measured current can be represented by equation 7.I(t) = I0 cos(ωt + φ) = Re(I∗ exp(iωt))I∗ = I′ + iI′′I0 = √I′2 + I′′2tan(φ) =I" I′Z∗ = Z′ + iZ" =U∗ I∗Where I′ is the real part of measured current, I′′ is the imaginary part of measured current, φ is the measured phase lag, Z′ is the measured real impedance of the material system, Z" is the imaginary part of impedance of the material system and the complex permittivity can be calculated byε∗(ω) = ε′ − iε" =−i ωZ∗(ω).1 C0Where C0 is the capacity of the empty sample capacitor.Recently BbDS has been used by several researchers [5-7] to characterize composite materials and were able to study the relation of dielectric properties to material performance. Raihan et, al.

Raihan et, al. [6] used Dielectric Relaxation Strength (DRS) which is the algebraic difference1539(6)(7)(8)(9)(10)(11)(12)between static permittivity value and the limiting frequency permittivity value as shown below in (13) to characterize the material.∆ℇ = ℇ𝑠 − ℇ∞Higher DRS indicates more interfacial polarization owing to charge accumulation aroundboundaries (voids, cracks etc.), indicating a lower mechanical strength.Design of ExperimentsIn the current study, the infill percentage and bed temperature were considered independent variables and the mechanical strength properties – peak force, stiffness, ultimate tensile strength (UTS) and dielectric property DRS are considered dependent variables. Table I shows different values considered for infill percentages and bed temperatures for the study.

Table I shows different values considered for infill percentages and bed temperatures for the study. For each case, 3 specimens were used to take in to consideration the statistics of each case.Table I Tabulated values of various infill percentages and bed temperaturesInfill Percentage (%) Bed-Temperature(oC)12.5 5033 7550 80100Specimen Preparation and Material SelectionDesignA 3D CAD model was designed using SolidWorks. For mechanical testing the samples were designed following the ASTM D638 [8] Type-1 standards as shown in Figure 4(a) with 7mm thickness and for dielectric testing circular samples of 40mm width with a thickness of 7mm were printed. The nature of the process leads to creation of rough surface because of which the electrode contact is not uniform with the sample surface.

The nature of the process leads to creation of rough surface because of which the electrode contact is not uniform with the sample surface. In order to improve the contact, two surfaces were marked 25x25mm square on either sides and two-part silver epoxy (working time 10 minutes and curing time 48 hours) was uniformly spread in these squares as shown in Figure 4(b).1540(13)Figure 4:- (a) Coupons for mechanical testing (b) Circular Specimens for dielectric testingMaterialPolylactic acid, one of the most common FDM materials is used for the current study. The greatest advantages for PLA in FDM printing are the facts that it creates no toxic gases while melting and so can be printed with no ventilation system, and it has low glass transition temperature.

It can be used to print with a variety of materials like, PLA, ABS, PVA, HIPS, Polyester (Tritan), PETT, bronze etc. Figure 5 shows the printer setup and samples which are being printed.1541(a) (b)Figure 5:- Specimen Preparation using Lulzbot Taz printer at UTARIDielectric Testing SetupAn in-house manufactured faraday cage was used to block out external interferences. The broadband dielectric/impedance spectrometer, manufactured by Novocontrol©, was used for carrying out the dielectric study. The sample was placed between two copper electrodes of 0.5- inch diameter embedded in blocks made of polycarbonate material. A voltage of 1.0 V is applied to the sample resulting in a flow of current with a phase shift. The dielectric properties of the material system are calculated using equations 10-12.

A voltage of 1.0 V is applied to the sample resulting in a flow of current with a phase shift. The dielectric properties of the material system are calculated using equations 10-12. Figure 6 shows the equipment setup that was used to carry out the experiments.Figure 6:- Dielectric Testing setup1542Mechanical Testing SetupAxial Tension test was performed on the MTS© tensile testing machine equipped with 50 KN load cell and hydraulic wedge grips at IPPM UTARI. The gripping pressure was approximately 12.5-150 psi (based on infill percentage) as shown in Table III. Static axial tensile loading was applied using displacement control at a constant rate of 0.025mm-s-1.

Static axial tensile loading was applied using displacement control at a constant rate of 0.025mm-s-1. The testing setup and a set of samples used for testing are shown in Figure 7.Table III MTS Testing Grip Pressures based on Infill PercentageInfill Percentage (%) 12.5% 33% 50% 100%Grip Pressure(psi) 12.5 50 100 150(a)(b)Figure 7:- (a) MTS setup and (b) Set of specimens used for testingResults & DiscussionA series of tensile tests were conducted on a set of AM fabricated samples with varyingprocess parameters. Figure 8 shows the fractured set of specimens.1543Figure 8:- Fractured set of specimens after mechanical testingThe observed mechanical and dielectric properties are tabulated below inTable IV. The ultimate tensile strength (UTS) and DRS values along with the coefficient of variation of the mechanical test data for each set is shown below.

This may be a result of material inflow across the voids due to high bed temperature. This additional material tends to reduce the internal residual stresses which ultimately increases the strength of the specimen. The permittivity as a function of frequency for various infills at different bed temperatures are shown in Figure 10.1545Figure 10:- Variation of real permittivity with frequency for various infill’s at different bed temperaturesFigure 11:- Comparison of mechanical and dielectric characterization for various infill’s at different bed temperatures1546From Figure 11, it can be observed that higher the DRS value, higher the ultimate tensile strength (UTS) in most of the cases with some discrepancies for 33 % and 50 % infill. This contradicts the observations made for polymer matrix composites by Raihan et,al. [5] which had different phases with different conductivity and permittivity where as these 3D printed coupons have only PLA filaments.

[5] which had different phases with different conductivity and permittivity where as these 3D printed coupons have only PLA filaments. This needs further analysis and will be published in subsequent journal publicationConclusion and Future WorkIn this work the influence of two independent build parameters on FDM manufactured specimens was investigated using mechanical and dielectric characterization of 3D printed samples. The investigated material was Polylactic acid (PLA). In most cases studied, higher DRS value corresponded to higher strength. It was observed that 50 °C bed temperature had the highest DRS and strength in most cases as it is lower than the glass transition temperature of the material system (60-65 °C), whereas the other bed temperatures used in the study (75,80 °C) are higher.Further investigation is required before a concrete conclusion can be made on the pattern observed here.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEffects of Centrifugal Disc Finishing for Surface Improvements in Additively Manufactured Gears Foxian Fan1*, Nicholas Soares2, Sagar Jalui2, Aaron Isaacson3, Aditya Savla1, Guha Manogharan2, and Tim Simpson1,21Department of Industrial and Manufacturing Engineering, The Pennsylvania State University, University Park 2Department of Mechanical Engineering, The Pennsylvania State University, University Park 3Applied Research Laboratory Gear Research Institute, The Pennsylvania State University, University ParkCorresponding author: fwf5049@psu.edu, The Pennsylvania State UniversityAbstract Additive Manufacturing (AM) is well suited to rapidly produce complex and customized geometries economically for low production runs. However, there is an inherent need for post- AM machining and surface finishing in most metal AM applications.

However, there is an inherent need for post- AM machining and surface finishing in most metal AM applications. Centrifugal Disc Finishing (CDF) is a media-based mass finishing process that can be employed to improve surface finish of external surfaces of AM parts with complex geometry. This original study aims to understand the influence of CDF processing conditions on Ti64 gear teeth fabricated via Powder Bed Fusion (PBF). A detailed statistical analysis is conducted to analyze the effectiveness of CDF to improve surface roughness of different build surfaces of the AM gear teeth. In addition, both contact profilometer and X-ray Computer Tomography (CT) techniques are applied to evaluate its effectiveness to measure CDF and AM surface finishing.

In addition, both contact profilometer and X-ray Computer Tomography (CT) techniques are applied to evaluate its effectiveness to measure CDF and AM surface finishing. Findings from this study on CDF of gear AM will benefit metal AM community by better understanding the impact of CDF processing conditions for surface improvements in mass finishing of metal AM parts.Keywords: Centrifugal disc finishing; AM surface finishing; AM gears; surface roughness; CT roughness measurements; powder bed fusionIntroduction In recent years, disruptive development in Additive Manufacturing (AM) has given manufacturers the ability to overcome previous geometric limitations in manufacturing processes. Commercial AM metal products, such as gears and other crucial mechanical components, have begun to be widely adapted to AM production.

Commercial AM metal products, such as gears and other crucial mechanical components, have begun to be widely adapted to AM production. Although AM offers many benefits in overcoming geometric limitations in different applications, higher geometry complexity has also increased the difficulty and cost in post-finishing procedures. Given this increase, the cost of post-finishing is becoming one of the largest factors in metal Additive1698Manufacturing cost. The National Institute of Standards and Technology (NIST) reported that the cost of AM post-processing can range from 4%-13% of the total manufacturing cost, depending on the process [1]. Similarly, Wohlers Report 2019 indicated that “more than 26% of the cost is from post-processing” [2]. Although typical post-processing procedures such as CNC machining provide good precision and surface finish, post-finishing procedures have increased lead time and cost due to low capacity, long setup time, and high machining cost.

Furthermore, CNC machining-based tool path planning and tool accessibility also have become a critical limitation in AM designs. As the geometry complexity in AM objects increases, post-processing cost and difficulties increase as well.Centrifugal Disc Finishing (CDF) is a common type of abrasive media-based mass finishing process for surface finish improvement. Similar mass finishing processes include vibratory bowl, drag finishing, stream finishing, and centrifugal bowl finishing. However, unlike many Mass Finishing processes, CDF does not require any fixtures or additional supports for parts. While the media-based mass finishing processes often require more time than similar machining processes, the setup time, capacity, and ability to handle complex parts can be superior to CNC machining.

The primary investigation in the paper is to determine the change in surface roughness of Powder Bed Fusion (PBF) built Ti-64 gear teeth after CDF post-processing, with respect to CDF process parameters.Literature Review The CDF machine consists of a stationary barrel with a disc shaped rotor at the bottom of the barrel (Figure 1). While running, the disc rotates at high speed and “the mass within the container is accelerated outward and then upward against the stationary side walls of the container, which act as a brake. The media and parts rise to the top of the load and then flow in toward the center and back down to the disc” [3]. The barrel size, media, and rotational speed vary depending on type of machine and part to be finished. CDF can be operated either with or without lubricant, known as wet or dry, respectively. Kitajima et al. [4] studied how the flow- through system in CDF affects the material removal rate and surface roughness.

Kitajima et al. [4] studied how the flow- through system in CDF affects the material removal rate and surface roughness. The lubricant type and flow rate impact were compared in terms of surface roughness and stock removal rate. The experiment finds the highest flow rate result in lowest surface roughness in comparison to lower flow rate and batched lubricant. Furthermore, stock removal rate is the highest when flow rate is set at lowest. According to user manual of Walther Trowal TT45 Centrifugal Disc Unit, key parameters of the CDF include media size, rotational speed, and lubricant concentration. Matsumoto et al. [5] investigated distribution of flow pressure of the media under both dry and wet conditions. The results showed that the bottom half of the barrel exhibited higher pressure under both conditions, and the stock removal in the dry processes was greater than that of the wet process.1699Figure 1.

Walther Trowal TT45 Centrifugal Disc Finishing UnitFurthermore, mechanical simulations have been conducted for centrifugal disc finishing. Sutowski et al. [6, 7] simulated the media flowing mechanism of centrifugal disc finishing. The result determined that the media in the bottom half of the barrel had higher velocity and energy. Cariapa et al. [8] developed a material removal rate prediction model for CDF with spherical ceramic media. The main predictors consist of the density ratio and hardness ratio of testing metal and media. The most appropriate exponential fitted model exhibited a maximum of 16% deviation. Although studies in CDF material removal rate or surface roughness prediction models are scarce, prediction models and for vibratory finishing processes have been studied through a few approaches including experiment based approach, surface geometry based approach, discreet element analysis, and computational fluid dynamics analysis, [9–12]. Vijayaraghavan et al.

Vijayaraghavan et al. [13] summarized literatures on varieties of mass finishing studies and conducted an experiment using a vibratory finishing process. A model for Ra (arithmetic mean) surface roughness based on media shape, conical and cylindrical, was developed and tested with input parameters consisting of orientation, lubricant concentration, and processing time.1700Jamal & Morgan [14] conducted an experiment to test surface roughness improvement effectiveness on Selective Laser Melting (SLM) fabricated Ti-64 samples. Subjects are tested through drag finishing, stream finishing, high energy centrifugal barrel finishing, and centrifugal disc finishing. Although the centrifugal disc was found to be the least efficient in comparison to the other processes, the author stated that the media type used in this study was not optimal. Boschetto et al. [15] conducted an investigation on using Centrifugal Barrel Finishing to process SLM fabricated Ti-64 parts.

Boschetto et al. [15] conducted an investigation on using Centrifugal Barrel Finishing to process SLM fabricated Ti-64 parts. This study considered stratification angle, rotational speed, and processing time as parameters that affected resultant surface roughness. The result indicated that 90-degree stratification angle specimens are harder to process than 0-degree specimens, while 45-degree specimens had medium behavior. Overall, there exist few studies in CFD application for post-processing AM parts. In this paper, the effect of CDF process parameters on the surface roughness of Ti-64 AM parts is examined and statistically analyzed to investigate applicability of CDF in metal AM post-processing.Methods Experiment Setup The 3D-printer used in this experiment was the 3D Systems Prox DMP 320 Powder Bed Fusion printer. The powder used was Ti6Al4V (Grade 23) with a nominal size of 15-45 micrometer.

The powder used was Ti6Al4V (Grade 23) with a nominal size of 15-45 micrometer. A total of 21 gear teeth samples were printed, each consisting of one and a half teeth (Figure 2). The nominal dimensions of the samples were 0.44 x 0.45 x 0.50 mm3 (Figure 3). The gear teeth were built vertically along the Z-direction, which was perpendicular to build plate, as seen in Figure 2. Each sample number was directly printed onto the structure using a boss on the top face (Figure 4). The support structure was manually scraped off the build plate and the samples did not undergo heat treatment or any other post-processing techniques before the CDF process.Figure 2. CAD model of the gear tooth samples used in this study, including build direction1701Figure 3. Dimensioned drawing of the gear teeth samplesFigure 4.

CAD model of the gear tooth samples used in this study, including build direction1701Figure 3. Dimensioned drawing of the gear teeth samplesFigure 4. Various views of the as-printed gear teeth samples1702CDF Machine Parameters The focus of this study was on investigating how rotational speed (RPM) and processing time (minutes) of the CDF process affect the surface roughness of the samples. The triangular ceramic media used is shown in Figure 5, with nominal dimensions of 4 mm x 10 mm. The loading ratio, lubricant flow rate, and finishing media remained fixed throughout the experiment. 18 of the samples were processed under six different conditions, in groups of three, and compared to the remaining three control (as-built) samples (Table 1).Figure 5. Triangular ceramic media used in CDF, with nominal dimensions of 4 mm x 10 mmTable 1.

Triangular ceramic media used in CDF, with nominal dimensions of 4 mm x 10 mmTable 1. CDF processing conditions for each gear teeth sample, including the processing speed (RPM) and duration (minutes).Sample Number 1, 2, 3 4, 5, 6 7, 8, 9 10, 11, 12 13, 14, 15 16, 17, 18 19, 20, 21CDF Processing Conditions 15 minutes, 200 RPM 30 minutes, 200 RPM 60 minutes, 200 RPM 30 minutes, 150 RPM 30 minutes, 250 RPM 30 minutes, 280 RPM NoneSurface Profilometer Data Acquisition A Taylor-Hobson Surtronic S128 profilometer was used to collect the surface roughness data for five different directions on each sample, shown in Figure 6. Directions 1 and 2 are along exterior face of the sample, parallel and perpendicular to the build direction, respectively. Directions 31703and 4 are along the interior tooth profile, parallel and perpendicular to the build direction, respectively. Direction 5 is along the exterior tooth profile, perpendicular to the build direction.

Direction 5 is along the exterior tooth profile, perpendicular to the build direction. While the most common roughness measurements in gear teeth analysis are taken along the tooth profile (directions 4 and 5), these five directions allow for the comparison of the CDF finishing capabilities in multiple areas. The first area of interest is the comparison of surface roughness for features parallel and perpendicular to the build layers. Next, the difference in roughness of external and internal features can be investigated. Finally, the surface profiles of flat faces can be compared to those of curved features. According to a review by Townsend et al. [16], ISO 4287 arithmetic mean (Ra) surface roughness is the most widely used surface metrology benchmark in AM literature.

According to a review by Townsend et al. [16], ISO 4287 arithmetic mean (Ra) surface roughness is the most widely used surface metrology benchmark in AM literature. In this study, both arithmetic mean surface roughness (Ra) and mean roughness depth (Rz) were evaluated statistically and determine the relationship between processing time, rotational speed, and surface roughness.Figure 6. CAD model of gear teeth sample, showing the five directions used for surface roughness measurementsSurface Inspection Using Micro-CT Micro-CT scans were obtained using a General Electric v|tome|x L300 nano/microCT scanner. In order to capture the data of small features, a voxel size of 11 µm was chosen. The raw scan data was imported into ImageJ. The stack was converted to 8-bit from 32-bit for ease of data handling.

In order to capture the data of small features, a voxel size of 11 µm was chosen. The raw scan data was imported into ImageJ. The stack was converted to 8-bit from 32-bit for ease of data handling. Processed image files were then transferred to Avizo for 3D image reconstruction.Discussion and Analysis Individual Surface Regression Results (Ra) The individual surface roughness data (Ra) collected for each group of test variable combinations can be seen in Figures 7-11. The Ra surface roughness linear regression result for1704Directions 1, 2, and 4 showed that rotational speed (RPM) was a significant parameter, at α=0.05 significance level (Figures A1, A2, and A4), but not significant for profiles 3 and 5 (Figures A3, A5). The parameter time (minutes) is not a significant parameter for all surfaces, using a α=0.05 significance level.

The parameter time (minutes) is not a significant parameter for all surfaces, using a α=0.05 significance level. The boxplots trends indicate external surfaces (Directions 1,2, and 5) showed more surface roughness reductions than internal surfaces (Direction 3,4). The summary of the statistical test results for each individual surface are summarized in Table 2. The Ra data from each surface is not consistent with the assumption that both processing time and rotational speed are significant parameters on resultant surface roughness. The authors assumed the reasons of statistical test failure can be attributed a few sources. A limited number of combinations of processing time and rotational speed were conducted. The initial surface roughness of the printed samples (unfinished) varied significantly. The phenomenon was noticed during the experiment while some as-build samples have much lower or higher surface roughness than average.Table 2.

The phenomenon was noticed during the experiment while some as-build samples have much lower or higher surface roughness than average.Table 2. ANOVA significance for Ra values of each surface profile, for both speed and duration and using a significance level of α=0.05Surface (Ra)RPM Significance (α=0.05) Time Significance (α=0.05)1SignificantInsignificant2SignificantInsignificant3InsignificantInsignificant4SignificantInsignificant5InsignificantInsignificantA categorical regression of all roughness data was then performed, taking into account the measurement location for each sample. The results indicated both surface and measurement location are significant parameters to the resultant surface roughens, at α=0.05 significance level. Processes duration (time) was found to not be a significant parameter in this regression test.1705Figure 7.

The linear regression results for Profiles 1 and 4 showed that rotational speed (RPM) was a significant parameter but process duration (minutes) was not significant, at α=0.05 significance level (Figures A7 and A10). Regression results for Directions 2 and 5 indicated both rotational speed (RPM) and duration (minutes) were significant parameters, at α=0.05 significance level (Figures A8 and A11). Neither rotational speed (RPM) nor duration (minutes) was significant at α=0.05 significance level for the Direction 3 (Figure A9). The analytical result of Rz is summarized in Table 3. Although not all individual surfaces regression were consistent with the assumption that both speed and time are significant parameter to resultant surface roughness, the boxplot of Rz showed a greater roughness reduction in comparison to the Ra plots. It also demonstrates the assumption that combination of the highest speed and duration resulted in the greatest decreases in surface roughness.

It also demonstrates the assumption that combination of the highest speed and duration resulted in the greatest decreases in surface roughness. Furthermore, the similar categorical regression with Rz indicated profile location, speed, and processing time are all significant parameters at α=0.05 significance level (Figure A12). The result proves that aside from speed and processing time, sample measurement location and direction had significant impact on the surface roughness. Additionally, while abrasive medias were able to reduce the external roughness gradually, the media size used in this experiment may have not been suitable for internal features, like surface 3, which showed least reduction and failed to satisfy all statistical tests. However, the Rz boxplot of surface 3 (Figure 14) did show a noticeable reduction in surface roughness in comparison to as-built samples (Samples 19-21).1708Table 3.

Rz values (µin) for each set of CDF processing conditions for interior tooth surface parallel to build direction (Direction 3)1710Figure 15. Rz values (µin) for each set of CDF processing conditions for interior tooth surface perpendicular to build direction (Direction 4)Figure 16. Rz values (µin) for each set of CDF processing conditions for exterior tooth surface parallel to build direction (Direction 5)The Rz regression analysis demonstrated a significant increase in the R-squared value in compression to Ra regression result. Furthermore, all predictor parameters were proved significant at 95% confidence level. One cause for the Rz statistical model being better than Ra1711model can be attributed to the processes in which CDF smooths the surface by removing “peaks” and leaving “valleys” mostly unaffected. A comparison of the trace profile for Sample 19 (control) and Sample 16 (280 RPM, 30 min) is shown in Figures 17 and 18.

A comparison of the trace profile for Sample 19 (control) and Sample 16 (280 RPM, 30 min) is shown in Figures 17 and 18. It can be seen the profilometer plots that the peaks on Sample 16 are much softer than the sharp ones seen on Sample 19. This same sentiment is corroborated by the CT reconstruction of the same samples in Figure 20. By observing the CT reconstructions, it can be seen that the CDF process dramatically effects the samples’ exterior edges by rounding the sharp corners. Therefore, Rz surface roughness which averages “peak” to “valley” distance on measured surfaces, seems more suitable for comparing this process. The statistical analysis and physical surface profile trace phenomenon demonstrated Rz is a better representation of surface roughness in comparison to Ra.Figure 17. Surface roughness profile for Sample 19 (control) as collected by profilometerFigure 18.

Surface roughness profile for Sample 19 (control) as collected by profilometerFigure 18. Surface roughness profile for Sample 16 (280 RPM, 30 min) as collected by profilometerFurthermore, significant lack of fit exists through all statistical analysis, which conclude that linear regression might not be suitable for this study. The CT scanned samples were1712reconstructed to visualize surface roughness change under different processing conditions. (Figure 19) The reconstructed images showed much more surface smoothing effect on the edges, while higher speed and longer time processing condition showed more aggressive surface smoothing effect on both surfaces and edges.1713200 RPMControl15 Min200 RPM200 RPM30 Min60 Min150 RPM280 RPM30 Min30 MinFigure 19.

CT reconstructions of select gear teeth samples after CDF processing1714Mechanical Test It is common knowledge in conventional metal manufacturing that fatigue strength can be enhanced by improving surface finish. To understand if the same principle is applicable for Additively Manufactured Ti6Al4V samples, Single Tooth Bending Fatigue (STBF) Gear Test (Modification of SAE J1619) was performed. Gear samples that finished under 200 RPM 15 min condition were Electro Discharge Machined(EDM) into single gear teeth. The testing setup is shown in Figure 20(a)&(b). The bending test machines used was 100 kN Servo-hydraulic Universal Test Machine with 30 Hz test frequency.Figure 20. STBF Test Setup: (a) Servo-hydraulic Bending Fatigue Test Machine (b) Gear Tooth in fixtureThree samples were tested. The first sample was tested with a maximum bending stress of 200 MPa and did not fail in 2,000,000 cycles.

The first sample was tested with a maximum bending stress of 200 MPa and did not fail in 2,000,000 cycles. The second sample was tested under 400 MPa bending stress and a crack was observed at 241,650 cycles on the root fillet on the opposite side of the tooth from the test fillet. This root is loaded in compression., see Figure 21. The same phenomenon occurred in Test 3 with the sample loaded to 500 MPa at approximately 250,000 cycles. In conventionally manufactured gears, the crack should initiate at the tensile side (the same side where load is applied). However, this unexpected behavior of crack initiation on compressive side calls for further metallurgical characterization. It is possible that the AM build contained porosity and/or intermetallic phases that are not typical in conventional gear materials. These defects could be susceptible to failure (crack initiation and growth) under compression that is not typical in wrought materials.1715Figure 21.

These defects could be susceptible to failure (crack initiation and growth) under compression that is not typical in wrought materials.1715Figure 21. Crack initiated on compressive side unexpectedly under 400 and 500 MPa.1716Conclusions & Future Work The overall regression analysis for Ra values showed the location of measurement and rotational speed were significant parameters, while processing duration was not significant, using a 95% confidence level. However, this can be attributed to limited variable combinations and the range of duration values used not being as wide as the set of rotational speed values. The overall regression for Rz showed that location of measurement, rotational speed, and processing duration were all significant predictors for surface roughness response, for the same confidence level. The profilometer profile trace showed significant reduction and rounding of sharp peaks.

The profilometer profile trace showed significant reduction and rounding of sharp peaks. CT image reconstruction of the gear samples also showed a noticeable rounding of sharp exterior corners. Although there was no direct relationship observed between the CT and profilometer data, it was determined that higher rotational speed and longer processing time resulted in lower surface roughness for both measurements. In consideration together with the regression result, Rz showed more roughness reduction and better statistical model reliability in comparison to Ra. However, linear regression showed high level of lack of fit in this study.The result has proved CDF applicability in surface improvement for AM external features. However, further work is needed to better optimization the process.

However, further work is needed to better optimization the process. This experiment only considered rotational speed and processing duration as predictors, while literature have shown other parameters such as loading ratio, media geometry/size, workpiece hardness, and lubricant flow rate to be significant in final surface roughness. For future work, further statistical analysis is needed to investigate more predictors and their interaction to determine an optimized material removal rate and surface roughness prediction model. Furthermore, geometric dimensions and tolerances need to be investigated for CDF processed AM parts so that the geometric dimensions and material removal relationship can be interpreted. Further investigation on AM gear material properties is also needed to explain the unexpected fatigue test behavior.Acknowledgement The authors acknowledge Penn State Gear Research Institute and Center for Quantitative Imaging for experiment preparation and data collection.

FECTS OF CRYOGENIC AGING ON A RAPID PROTOTYPED (RP) POLYMERF. Fuentes1, A. Gallardo Jr1., J. Martinez1, J. Bulman1, J. Foyos1,M. Mendelson1, R. Noorani1, B. Fritz21NSF Research Experiences for the Undergraduates (REU) Loyola Marymount University, Los Angeles, CA - 900452Northrop Grumman , El Segundo, CA - 91355ABSTRACTLittle research has been done on the post-processing (aging) of rapid prototyped (RP) polymers at temperatures below 123K (–238˚F). Test specimens of RP thermosetting resin (DSM-Somos 8110) were fabricated and cryogenically aged from 10-25 hours. The tensile strength and impact toughness were measured. This work will study the effect of cryogenic aging on yield strength of Somos 8110. This paper will also discuss our interpretation of the data based on fractography.INTRODUCTIONAs industry continues to increase its use of rapid prototyping (RP) materials, the necessity for enhancing the properties of RP materials is becoming more important.

Past studies have concluded that by reducing the temperature below 123K (–238˚F), cryogenic processing increased the strength of many materials, particularly metals. But while the strength of the materials increased, the cryogenic treatments also reduced ductility.Though relatively unstudied, the results of cryogenic tempering on certain polymers will prove highly valuable to industry. This study hopes to increase the usage of RP materials through cryogenic processing by optimizing the combination of strength and ductility for the stereolithography photopolymer epoxy resin, DSM Somos 8110 [1].The cryogenic test procedure began with the construction of the tensile and impact test samples using the RP materials. In order to increase the mechanical properties, these samples were then subjected to 88K (-300˚F) using several different holding and ramping times. Next, the samples were examined using tensile and impact tests. Finally, the data obtained from the tests was analyzed.

Next, the samples were examined using tensile and impact tests. Finally, the data obtained from the tests was analyzed. Based on metals, it was expected that the tensile strength of the RP materials would increase, and the toughness would decrease after cryogenic processing. However, we hoped there would be a combination of cryogenic aging treatments that would improve both the strength and toughness.The purpose of this paper is to discuss the cryogenic method for post-processing rapid prototyped DSM-Somos 8110 epoxy resin. In addition, this paper will examine the effect of cryogenic aging on the tensile strength and impact energy.115BACKGROUNDCryogenicsCryogenics is the science and art of producing a low temperature environment. It started in 1877 when two scientists, Cailletet in Paris and Pictet in Geneva, developed a procedure to liquefy oxygen in the laboratory [2]. Now, liquid nitrogen is one of the most common cooling media.

Now, liquid nitrogen is one of the most common cooling media. Since the normal boiling points of nitrogen and other permanent gases such as helium, oxygen, and argon is about 84.8K (≈-307˚F), the cryogenic temperature is generally considered to be 84.8K or below [2].Cryogenic processing is an important field in industry today. For example, an increase of 195% to 817% in wear resistance has been reported for standard steel that was cryogenically treated [3]. The cryogenic process consists of three stages that are time and temperature dependent to the aging temperature. This process starts with the gradual ramping down of temperature for a pre-determined time range, to a specific point. The temperature is then held for a period of time. Finally, a gradual process of temperature ramp-up is executed.

The temperature is then held for a period of time. Finally, a gradual process of temperature ramp-up is executed. As a result of cryogenic processing, some unknown molecular changes occur, modifying the structure of the material [3].Conventionally, heat treating is used to strengthen steel and other alloys. Upon quenching steel below its eutectoid decomposition temperature, some retained austenite remains [3]. However, cryogenic processing transforms virtually all of the retained austenite to martensite, along with a precipitation of fine carbide particles. These effects have resulted in increased wear resistance [3].Polymers and StressesPolymers can be classified according to their chain structure and chain arrangement. Chain structure of polymers is further characterized by the chemical composition of monomers and their relative arrangement within a chain. The common chain arrangements are linear chains, branched chains, and cross-linked networks (e.g.

The common chain arrangements are linear chains, branched chains, and cross-linked networks (e.g. epoxy resins).The general characterization of polymers is very difficult due to the complicated structures, the great variety of chemical compositions and the transfer of molecular groups [4]. However, the fracture strength of polymers can be explained from a microstructural point of view. These factors can include molecular stress concentrations and residual thermal stresses [5].The twisting and kinking of polymer chains contributes to local stress concentrations due to the uneven stress distribution between the molecules. When the entangled polymer chains are under stress, the chain segments in the load direction provide most of the load-carrying capacity.

When the entangled polymer chains are under stress, the chain segments in the load direction provide most of the load-carrying capacity. When polymers are cooled too quickly during cryogenic processing, thermal stresses are built up within the polymer, and these act to reduce the tensile strength [4].116EXPERIMENTAL PROCEDUREEquipment and ProcessThe following major equipment was used in this research.1. Northrop Grumman SLA-250 RP Machine 2. Cryogenic Treatment Equipment 3. Instron Tensile Testing Machine 4. Izod Impact Tester 5. Scanning Electron Microscope (SEM) The experimental setup if shown in Figure 1.Temp ControlCryogenic ChamberVacuum PumpCryogenic PumpFigure 1. Experimental setup for cryogenic processing.The following outlines the steps used in the experimental process that was used to design,fabricate, test and analyze the samples.1. Design Dog Bone and V-Notch Samples 2. RP Machine Builds Parts from Design 3.

Design Dog Bone and V-Notch Samples 2. RP Machine Builds Parts from Design 3. Expose Parts to Cryogenic Treatment with Liquid He. 4. Tensile and Impact Testing 5. SEM FractographyDesigning and Prototyping the SamplesDrawings of the dog bone and v-notch shaped samples were created using AutoCAD (Figs. 2a and 2b) and these were saved as separate .DWG files. These files were then converted into STL format for use with stereolithography software. SLA-250 (Northrop Grumman) RP machines was then used to rapid prototype the parts.117Figure 2a. Two dog bone-shaped samples.Figure 2b. Two v-notch-shaped samples.Cryogenic TreatmentThe following process was used to cryogenically treat each sample.1. All samples except the baseline went through cryogenic treatment before testing. The samples were prepared at Northrop Grumman, and cryogenically aged at Loyola Marymount University.2.

All samples except the baseline went through cryogenic treatment before testing. The samples were prepared at Northrop Grumman, and cryogenically aged at Loyola Marymount University.2. The cryogenic process is characterized by three parameters: ramp-down time from room temperature to 88K (–300oF), hold time at 88K, and ramp-up time from 88K to room temperature.3. Preliminary experiments were performed on cryogenic treatment. The ramp-down time of 14 hours was used. Three holding times of 10, 15 and 20 hours, and ramp up times of 14 hours were used.4. Samples are labeled as follows: XX-XX-XX. The numbers represent the ramp-down time, holding time and ramp up time, respectively, in hours. (e.g: 14-10-14 means that the samples were ramped-down in 14 hours, cryogenically aged 10 hours, and then ramped-up in 14 hours.) Tensile TestingThe yield and ultimate strengths of the samples were measured using the Instron Universal Testing Instrument 4500.

The cross-head speed of the test machine was 0.0212 mm/s. The interface with the machine was performed using the front panel and a software program running on a desktop computer. The test specimens were designed in the shape of a dog-bone, similar to ASTM D638-97. The end pads of the specimen were tightened and aligned between the two grips of the frame unit. A computerized load cell located inside the frame unit measured the force applied to the gauge section of the sample. The software calculated the tensile stress in the gauge section. Strain was measured with an extensometer, and the stress vs. strain curves were plotted as the specimen was continuously loaded. The yield stress was measured at 0.2% offset strain, and the ultimate strength was measured at the maximum stress the material could withstand. Four to six specimens were tested for each cryogenic aging treatment (0-20 hours).

Four to six specimens were tested for each cryogenic aging treatment (0-20 hours). The tensile strength data were statistically analyzed to determine the effect of aging treatment on the yield and ultimate strengths.Izod Impact TestingIzod impact testing is performed to determine the toughness of a material. The sample is made118with a centered v-shaped notch. During the impact testing, the sample is subjected to a quick and intense blow by a hammer pendulum. The impact test evaluates the material’s resistance to crack propagation. The impact energy absorbed by the sample during failure is determined by calculating the difference in potential energy of the hammer [5].SEM FractographyThe fracture surfaces of the failed tensile samples of Somos 8110 resin were examined under a scanning electron microscope (SEM). SEM fractography was used to interpret the results of the tensile strength data for both the untreated (baseline) and cryogenically treated samples [6].

SEM fractography was used to interpret the results of the tensile strength data for both the untreated (baseline) and cryogenically treated samples [6]. The SEM used an acceleration voltage of 1.2kV. The fracture surfaces were observed sequentially at magnifications of 15x, 40x and 150x.The stretching of the polymer chains and fracture initiation were correlated with the yield strength data. Here critical-size cracks were initiated from defects and propagated when the yield strength of the polymer was exceeded. The propagation of the crack front was correlated with the ultimate strength. Then the crack propagation patterns could be observed on the fracture surface of the failed specimens.RESULTS AND DISCUSSIONTensile and Impact TestingThe results of the yield strength and ultimate strength vs. cryogenic treatment time (0-20 hours) are shown in Figures 3A, 3B. In both cases, the strengths appear to be affected by aging time.

cryogenic treatment time (0-20 hours) are shown in Figures 3A, 3B. In both cases, the strengths appear to be affected by aging time. To verify these results, the data were statistically analyzed using multiple t-testing, which compared the means of two treatments at-a-time at a 0.05 level of significance [7]. A one-way analysis of variance (ANOVA) could not be used, because the variances of the treatments were unequal. For the yield strength data (Figure 3A), it was determined with 95% confidence that the means of the treatments were statistically insignificant. Due to the large variances at 15 and 20 hours, we can say that there was no significant increase in the yield strength with aging time.For the ultimate strength (Figure 3B) vs. aging time, the data were analyzed in the same way. Again, one-way ANOVA could be not be used.

aging time, the data were analyzed in the same way. Again, one-way ANOVA could be not be used. Multiple t-testing on the treatment means showed that for 95% confidence two conclusions were reached: (1) the mean strengths at 0, 10 and 20 hours were equivalent, and (2) the mean strength at 15 hours was significantly lower than that at 10 and 20 hours. Hence, aging time produced a drop in the ultimate strength between 10 and 15 hours. Otherwise, there was no significant effect of cryogenic aging on the ultimate stress.When the impact energy data vs. aging time (Figure 3C) were statistically analyzed, the large variance at 10 hours indicated that the treatment means were insignificant with 95% confidence. Hence, there was no effect of cryogenic treatment time on the impact energy.119SEM FractographySince all of the tensile specimens failed near the fillet radius, failure analysis was performed on the tested specimens.

The fracture surfaces of the tensile specimens were examined under three different treatment conditions: (1) Sample A, treated for 15 hours and having the highest yield strength (Figure 3A), (2) Sample B, treated for 15 hours and having the lowest yield strength (Figure 3A) and (3) Sample C, untreated.The following findings were made from observing the fracture surfaces of Samples A, B, C in Figures 4A, 4B and 4C, respectively. In all cases, the fracture originated from surface defects on the fillet radius (between the gauge length and grips) of the tensile specimens. Fracture propagated from left to right, as shown by the 'river markings' on the fracture surface [6]. Samples B and C showed similar fracture patterns that consisted of mixed mode I and III fracture [6]. Here, fracture initiated from the stepped fillet radius in Figures 4B, 4C and 5B, which was formed during stereo-lithographic processing.

Here, fracture initiated from the stepped fillet radius in Figures 4B, 4C and 5B, which was formed during stereo-lithographic processing. Sample A exhibited primarily mode I fracture from the fillet surface, and fracture did not originate from steps on the fillet radius (Figures 4A and 5A).Samples B and C had low yield strengths and similar fracture patterns. The low strengths were interpreted as being caused by the stepped fillet radius. The steps probably caused a high stress concentration at the fillet surface. In addition, the steps appeared to have surface defects on the fillet radius (Figures 4B, 4C and 5B). Sample A had the highest yield strength, and there were no steps or surface defects on the fillet radius where fracture originated. Hence, the stepped fillet surfaces with their associated surface defects probably caused the specimens to fail prematurely at lower strengths.

Hence, the stepped fillet surfaces with their associated surface defects probably caused the specimens to fail prematurely at lower strengths. The steps and surface defects in Samples B and C also accounted for the higher variance (scatter) in the data.The differences in the fracture patterns between Sample A and Samples B, C could have also accounted for their difference in yield strength. In Samples B and C, it appeared that the crack front propagated at slower velocities than in Sample A. The crack had more time to separate on to different levels and develop mode III shear lips (Figures 4B and 4C). In Sample A, there were no steps and surface defects at the fillet radius. Consequently, a critical-size crack had to be nucleated on the fillet surface, and this required extra stress, which caused the strength to be higher. Once a critical-size crack was formed, the extra stress caused the crack to propagate catastrophically at a higher velocity.

Ultimate Strength vs. Cryogenic Aging Time.Figure 4A. 40x SEM micrograph of Sample A in Figure 3A.Figure 4C. 40x SEM micrograph of Sample C (Baseline) in Figure 3A.25Figure 5A. 15x SEM micrograph of Sample A from Figure 4A.Figure 5B. 15x SEM micrograph of Sample C (Baseline) from Figure 4C.CONCLUSIONS AND RECOMMENDATIONSBased on the findings from our research, the following conclusions and recommendationscan be made: 1. DSL-Somos 8110 test specimens were fabricated by stereo-lithography using SLA-250 RP equipment. The specimens were laser cured by pulling the specimens parallel to the tensile axis. 2. The test specimens were cryogenically treated by a ramp-down cycle from room temperature to 88K in 14 hours, a hold cycle of 10 - 20 hours, and then a ramp-up cycle from 88K to room temperature in 14 hours. 3. Due to large data scatter, the yield strength and impact energy were not affected by cryogenic aging treatment.

3. Due to large data scatter, the yield strength and impact energy were not affected by cryogenic aging treatment. Only the ultimate strength exhibited a significant decrease with aging treatment from 10 to 15 hours. 4. It was expected the wide data scatter was caused by microscopic steps and their associated defects at the fillet radius of the tensile specimens. Samples with steps and defects had lower strengths, and those without steps and flaws had high strengths. The variability in the surface texture of the fillet radius caused a wide variability in the tensile strength and impact energy. 5. For future work, it is recommended that the SLA specimens be rapidly prototyped normal to the face of the dog-bone specimen. This will hopefully reduce the data scatter, which currently masks any potential effect of cryogenic aging time on the tensile strength and impact energy.ACKNOWLEDGEMENTSThe work was funded by an NSF Grant under Research Experiences for Undergraduates (REU).

The authors also appreciate the support provided by the Northrop Grumman Corporation. We also acknowledge the contributions toward the development of this paper by our students, J.C. Jackson and Joey Weinmann.122REFERENCES1. 9, and Q79.Stratasys, 1996, “FDM System Documentation”. F1650-5, F1650-6, F1650-11, G-8, G-2. University of Southampton, UK. Oxford: Clarendon Press.Scurlock, R.G., 1992, “History and Origins of Cryogenics,” Institute of Cryogenics,3. Future to Our Fingertips,” Business, July, 1994.Chhabra, V., “Ingenious Cryogenics: Gainsville Inventor Frank Masyada Brings the4. New York and London.Hartwig, G., “Polymer Properties at Room and Cryogenic Temperatures,” Plenum Press,5. Publishing Company, BostonAskeland, D.R., 1995, “The Science and Engineering of Materials,” Third Edition, PWS6.

FECTS OF CRYOGENIC PROCESSING ON RAPID PROTOTYPING MATERIALS (DSM SOMOS-8110 AND DURAFORM PA)J.Jackson1, G. Chapple1, J. Do1, X. Zhuang1, J. Bulman1, J. Foyos1, M. Mendelson1, R. Noorani1, B. Fritz21NSF Research Experiences for the Undergraduates (REU) Loyola Marymount University, Los Angeles, CA - 900452Northrop Grumman , El Segundo, CA - 91355ABSTRACTThis research investigates the effects of cryogenic processing on the properties of rapid prototyped materials. Not much research has been done on the post-processing (aging) of rapid prototyped (RP) polymers at temperatures below 159K (–173˚F). Test specimens of RP thermoplastic resin DSM-Somos 8110 and DuraformPA Nylon were fabricated and cryogenically aged from 5-30 hours. The tensile strength and impact toughness were measured. The goal of this work was to study the effect of cryogenic aging on yield strength and ductility.

The tensile strength and impact toughness were measured. The goal of this work was to study the effect of cryogenic aging on yield strength and ductility. This research investigated (1) the cryogenic aging of DSM-Somos 8110 and DuraformPA Nylon, (2) the effects of controlled ramp-downs/ups on the ultimate and tensile strengths of samples, (3) the experimental methods, and (4) the analysis and interpretation of the data.INTRODUCTIONThe primary objective of this research project is to apply cryogenic processing to Rapid Prototyped materials. Cryogenics is the science and art of producing cold. It started in 1877 when two scientists, Cailletet in Paris and Picet in Geneva, developed a procedure to liquefy oxygen in a laboratory [1-2]. Nowadays, nitrogen and helium are the most common cooling media. Since the normal boiling points of nitrogen and other permanent gases such as helium, oxygen and argon are about 120 K (approx.

Nowadays, nitrogen and helium are the most common cooling media. Since the normal boiling points of nitrogen and other permanent gases such as helium, oxygen and argon are about 120 K (approx. –244 oF), the cryogenic temperature is generally considered 120 K or below [3-4].Cryogenic processing is one of the most important fields in industry today [4]. It helps to reduce costs for industry and increase industrial efficiency. For example, industrial application has reported 195% to 817% increase of wear resistance for standard steel that was cryogenically treated [2]. The cryogenic process consists of three stages based on time and temperature variables. This process starts with gradual ramping down of temperature to a specific point, the temperature is held at that point for a period of time, then the temperature is brought up to room temperature.

This process starts with gradual ramping down of temperature to a specific point, the temperature is held at that point for a period of time, then the temperature is brought up to room temperature. As a result of this deep cooling and heating cycle, molecular changes occur, binding the atoms in the metal together [4].Over the last 10-20 years, relatively little research has been conducted on cryogenic processing. Most research in cryogenics has been preformed on metals. Rapid Prototyping (RP) is a new technology the takes information from a computer-aided design file and makes a 3D part by building it one layer at a time [5]. When RP was first introduced in the late 80’s, the607materials used to produce the parts had low yield strength. This experiment attempts to show that the strengths of RP materials can be increased by cryogenically processing these parts before industrial application.

This experiment attempts to show that the strengths of RP materials can be increased by cryogenically processing these parts before industrial application. As mentioned before, in all the cryogenic work the scientists and engineers lowered the temperature of the sample to cryogenic condition very fast, held it to the temperature for a few hours and then ramped up the temperature as fast as possible [4]. Two years ago, the authors did research on the cryogenic processing of both ABS plastic using Fused Deposition Modeling (FDM) and DSM SOMOS 8110 from Stereolithography Apparatus (SLA)- 250 machines. While the cryogenic processing did not have much effect on ABS plastic, the yield strength of DSM SOMOS was increased between 25-50% [6].

While the cryogenic processing did not have much effect on ABS plastic, the yield strength of DSM SOMOS was increased between 25-50% [6]. The authors believe the ramp-down and ramp–up conditions might make a significant contribution to the increase in strength of the polymer samples.The authors have recently developed a Data Acquisition System (DAS) using LabVIEW from National Instruments that provided the programming power for the ramp-down and ramp- up conditions. The conventional and proposed cryogenic processes are shown in Figures 1 and 2.EXPERIMENTAL PROCEDUREEquipment and ProcessThe following major components were used in this experiment:1. Northrop Grumman SLA-250 RP Machine 2. Cryogenic Treatment Equipment, Figure 3 3. Instron Tensile Testing Machine 4. Izod Impact Tester 5. Scanning Electron Microscope (SEM)608Temp ControlCryogenic ChamberVacuum PumpCryogenic PumpFigure 3.

Instron Tensile Testing Machine 4. Izod Impact Tester 5. Scanning Electron Microscope (SEM)608Temp ControlCryogenic ChamberVacuum PumpCryogenic PumpFigure 3. Cryogenic EquipmentThe following Figure 4 outlines the experimental process that was used to design, fabricate,test, and analyze the samples.Design Dog Bone and V-Notch SamplesRP Machine Builds Parts from DesignTensile and Impact TestingExpose Parts to Cryogenic TreatmentFigure 4. Outlines of Steps for the ExperimentDesigning and Prototyping the SamplesDrawings of the dog bone and v-notch shaped samples were created using AutoCAD (Figs. 5A and 5B) and then were saved as separate .DWG files. These files were then converted into .STL format for use with QuickSlice software. The SLA-250 (Northrop Grumman) RP machine was then used to rapid prototype the parts.609Figure 5A. Two dog bone-shaped samples. Figure 5B.

The SLA-250 (Northrop Grumman) RP machine was then used to rapid prototype the parts.609Figure 5A. Two dog bone-shaped samples. Figure 5B. Two v-notch-shaped samples.Cryogenic TreatmentThe following procedure was used to cryogenically treat each sample.1. All samples except the baseline went through cryogenic treatment before testing. The samples were prepared at Northrop Grumman and processed at Loyola Marymount University.2. The cryogenic process is characterized by three parameters: ramping down time from room temperature to 82K (–312oF), holding time at 82K, and ramping up time from 82K to room temperature.3. Preliminary experiments are then performed on the samples with the ramping cryogenic treatment process. Ramp-down times of 5 hours are used. Holding times of 5, 8, 10, 12 and 20 hours, and ramp up times of 5 hours are used.4. Samples are labeled as follows: XX-XX-XX.

Ramp-down times of 5 hours are used. Holding times of 5, 8, 10, 12 and 20 hours, and ramp up times of 5 hours are used.4. Samples are labeled as follows: XX-XX-XX. The numbers represent the ramp-down time, holding time and ramp-up time, respectively, in hours. (e.g: 05-10-05 means that the samples were ramped-down in 5 hours, held in a cryogenic state for 10 hours, and then ramped-up in 5 hours.)Tensile TestingThe tensile strength of the samples was measured using the Instron Universal Testing Instrument 4500. The cross-head speed of the test machine was 0.0212 mm/s. The machine was interfaced using its front panel and a software program running on a desktop computer. The test specimens were created in the shape of dog-bones, as shown in Figure 5A. The dog-bone shaped samples were held in place by the two opposing grips of the testing platform. A computerized load cell located inside the frame unit measured the force applied to the dog-bone shaped samples.

A computerized load cell located inside the frame unit measured the force applied to the dog-bone shaped samples. Strain was measured with an extensometer, and the stress vs. strain curves were plotted while the specimen was gradually loaded. The yield stress was measured at 0.2% offset strain, and the ultimate strength was measured at the maximum stress the material could withstand.Four to six samples were tested for each cryogenic aging treatment (0-20 hours). The tensile strength data was statistically analyzed to determine the effects of aging treatment on yield and ultimate strength.610Izod Impact TestingIzod impact testing was performed to determine the toughness of the treated material. The samples were made with a centered v-shaped notch. During the impact testing, the samples were subjected to an impulsive blow by a hammer pendulum.

The samples were made with a centered v-shaped notch. During the impact testing, the samples were subjected to an impulsive blow by a hammer pendulum. The impact test evaluates the material’s resistance to crack propagation.RESULTS AND DISCUSSIONTensile and Impact TestingThe results of the yield strength, ultimate strength and impact energy per unit area vs. cryogenic treatment time (0-30 hours) are shown in Figures 6A, 6B, 6C, 6D, 6E and 6F. For the yield strength of the Somos 8110, it appears to peak at a hold time of 10 hours, which is consistent with the results of the previous year, though the results show too much variance to determine a definite trend. Three of the four points of the 10 hour hold time are above the base line, so an increase is inferred, Figure 6A. On the other hand, the 10 hour hold time appears to be a low point of the ultimate tensile strength graph for the Somos 8110.

On the other hand, the 10 hour hold time appears to be a low point of the ultimate tensile strength graph for the Somos 8110. Again, the scatter of data points makes it difficult to discern a trend, Figure 6B.Yield StrengthUltimate Tensile Strength112410229) a P M20) a P M( h t g n e r t S d e Yli876( h t g n e r t S el is n e T e t a m1816145i t lU4123100246810121402468101214Hold Time (Hours)Hold TIme (Hours)Experimental DataAverageExperimentalAverageFigure 6A. Somos 8110 Yield Strength.Figure 6B. Somos 8110 Ultimate Strength.The impact energy of the Somos 8110 appears to decrease as the holding time increases. The trend, again, is complicated to find because of the erratic nature of the data -- even the baseline samples have a large degree of scatter, Figure 6C.

The trend, again, is complicated to find because of the erratic nature of the data -- even the baseline samples have a large degree of scatter, Figure 6C. As for the yield strength of the DuraformPA nylon, there is a significant drop for the hold time of 12 hours, Figure 6D. Further testing will need to be done to determine a reason for this effect.611Impact EnergyYield Strength2.52423222) J ( h t g n e r t St c a p m1.51) a P M( h t g n e r t S d l e i Y212019180.51716015024681012140510152025Hold Time (Hours)Hold Time (Hours)Experimental DataAverageExperimental DataAverageFigure 6C. Somos 8110 Impact Energy.Figure 6D. DuraformPA Yield Strength.The ultimate strength appears to increase with the increase in hold time, Figure 6E. Once again there is a significant scatter of data and the true trend is unclear.

DuraformPA Yield Strength.The ultimate strength appears to increase with the increase in hold time, Figure 6E. Once again there is a significant scatter of data and the true trend is unclear. There is a significant increase in the impact energy at the 5 hour hold time, Figure 6F, with a variance of the data points as in the previous samples. The v-notch samples varied based on their position in the copper carrier. The v-notch samples were arranged four in a row. The middle samples had lower impact energies than the outside samples.Ultimate Tensile StrengthImpact Energy472461.6) a P M( h t g n e r t S e l i s n e T4544) J (y g r e n Et c a p m1.20.8430.4420051015202505101520Hold Time (Hours)Hold Time (Hours)Experimental DataAverageExperimental DataAverageFigure 6E. Duraform Ultimate Strength.Figure 6F.

Duraform Ultimate Strength.Figure 6F. Duraform Impact Energy.61225The effects of ramp-down times on the yield strength and ultimate strength are shown inFigures 7A and 7B. In both cases, the strengths appear to be negatively affected.FractographyThe DuraformPA Nylon samples were too ductile to produce a flat fracture surface. The plasticity was demonstrated by the necking of the sample as shown in Figure 8A, which also illustrates the rough surface of the break.The Somos 8110 samples broke cleanly, producing a clear mirror which can be seen in Figure 8B. All Somos 8110 samples contained an array of dimples along one side, which were a byproduct of the rapid prototyping process. The dimples introduced weak points and can be seen in Figure 8C. The parts always broke along the dimples and this can be seen in Figure 8D.Figure 8A. Duraform PA Nylon Necking Figure 8B. 36x SEM micrograph ofDSM-Somos 8110613Figure 8C. Top view of DSM-Somos Figure 8D.

Duraform PA Nylon Necking Figure 8B. 36x SEM micrograph ofDSM-Somos 8110613Figure 8C. Top view of DSM-Somos Figure 8D. 36x Somos 8110 Dimples andfracture surfaceCONCLUSIONS AND RECOMMENDATIONSBased onthe findings fromthis experiment,the following conclusions andrecommendations can be made:1. Due to large data scatter it is difficult to determine the trend of the data. 2. The defects in the Somos 8110 samples from the previous year were corrected resulting in improved data, but the Somos 8110 contained an array of dimples that created weak points. If these dimples could be removed from the sample the results may improve. 3. The v-notch parts of both the Somos 8110 and the Duraform 8110 should be arranged around the perimeter of the copper carrier of the cryogenic chamber to ensure a uniform heat distribution to each sample during the ramp-up phase.4.

Cryogenic processing of prototypes samples has not significantly and consistently impacted the strengths of the samples. As the strengths of the rapid prototyping materials are increased through research and development, cryogenic processing of rapid prototyped samples may not be necessary. ACKNOWLEDGEMENTSPart of this work was funded by a NSF Grant under Research Experiences for Undergraduates. The authors also appreciate the support provided by the Northrop Grumman Corporation, as well as Dr. Bulman and Joe Foyos for their support in testing the materials.6141.2.3.4.5.6.REFERENCESScurlock, R.G., 1992, “History and Origins of Cryogenics,” Institute of Cryogenics, University of Southampton, UK.

fects of Cu and SiO2 on Laser Sintering of PolycarbonateK.M. Fan, W.L.CheungDepartment of Mechanical Engineering, The University of Hong Kong, Pokfulam Road, Hong Kong, China.AbstractAdditives of different thermal properties, Cu and SiO2 (amorphous and crystalline), were blended to polycarbonate (PC) powder to modify its heat transfer properties and fusion behavior during laser sintering. The blends were sintered under different energy densities of the laser beam to produce mono-layer films. The surface morphology and the thickness of the films were studied. When sintering under the same condition, the composite films which contained a high content of Cu powder exhibited a more porous surface structure. This was caused by the increased heat loss because the Cu powder has a higher thermal conductivity. Also, the solid Cu particles would hinder the flow of the molten polymer, resulting in a low degree of fusion.

Also, the solid Cu particles would hinder the flow of the molten polymer, resulting in a low degree of fusion. For a given Cu powder content, reducing its particle size gave a more porous surface structure and a smaller thickness of the sintered films. This was probably due to a more even distribution of the fine Cu particles, which increased the heat loss and reduced the effective amount of energy for fusion. On the other hand, fine SiO2 and quartz powders caused degradation of the polymer because of the increased energy dissipation near the film surface and poor heat transfer properties of the additives comparing with Cu.Keywords: Laser sintering; PC/Cu; PC/SiO2; fusion behavior.1. IntroductionIn laser sintering processes, the energy source is very intense and the time for fusion is extremely short. In addition, polymer materials are poor thermal conductors.

IntroductionIn laser sintering processes, the energy source is very intense and the time for fusion is extremely short. In addition, polymer materials are poor thermal conductors. Under normal sintering conditions, the polymer particles are only slightly fused together thus the sintered components are porous and have a coarse surface finish. This not only affects their mechanical strength but also the dimensional accuracy. One possible way of improving the fusion behavior of the polymer powder is by modifying the thermal properties of the material. Polycarbonate was blended with Cu, quartz and amorphous SiO2 and the blends were sintered using a laser- engraving machine.Composite materials are not rare in the SLS process, e.g. DuraForm Glass Filled, Cu with polyamide (DTM products), and Cu with PMMA[1]. The role of glass spheres in DuraForm is to increase the strength and reduce in the shrinkage of the sintered components.

DuraForm Glass Filled, Cu with polyamide (DTM products), and Cu with PMMA[1]. The role of glass spheres in DuraForm is to increase the strength and reduce in the shrinkage of the sintered components. In the last two material systems, Cu forms the main body of the green parts and the polymers only serve as a binder. In this work, copper, SiO2 and quartz were added to change the thermal properties of the material systems. The effects of content and particle size of the additives on surface morphology and thickness of the sintered films were studied. This paper outlines some preliminary results of the work.1102. Materials and processingThe PC powder was supplied by DTM, trade name LPC-3000. It has a glass transition temperature of 150(cid:176) C. The Cu powder was ground from the Cu infiltrant by DTM. The amorphous SiO2 and quartz powders were products of Fisher Scientific. The original size range of quartz was from 200 to 800 mm.

The Cu powder was ground from the Cu infiltrant by DTM. The amorphous SiO2 and quartz powders were products of Fisher Scientific. The original size range of quartz was from 200 to 800 mm. After grinding, the average powder size reduced to 140 mm. Figure 1 shows the PC powder and the additives before sintering, and table 1 lists some of their physical and thermal properties.a) PC powderb) Cu powderc) SiO2 powderd) Quartz powderFigure 1. SEM micrographs of polycarbonate, Cu, SiO2 and quartz powders.Table 1. Physical and thermal properties of the polycarbonate and additives [2,3].MaterialPCCuamorphous SiO2Quartzair (at 300K)Thermal conductivity (W/(m.k)) Specific Heat Capacity (J/Kg.K) Average particle size (mm) Density (g/cm3)0.181206900.6240038528 & 2936.74322.568.147951402.640.02631046111Sintering was carried out on a M25e Universal Laser Engraving Machine.

According to Nelson[4], the energy density (ED) of the laser beam during selective laser sintering (SLS) can be expressed as follows:P 3 fED =BS 3 SCSPwhere P is fill laser power, BS is beam speed, SCSP is the scan spacing and f is a conversion factor. In this work, ED was varied by different settings of P and BS. An increase in P or decrease in BS would result in a higher ED and vice versa. The number of pulses per inch (PPI) of the laser beam along the scanning direction and the number of dots per inch (DPI) perpendicular to the scanning direction were kept constant at 500. The sintered samples were examined with a Cambridge Stereoscan 360 scanning electronic microscope (SEM) at an operating voltage = 20kV and working distance = 20mm. The thicknesses of the films were measured from SEM micrographs of the cross sections and verified by a caliper. The samples were coated with gold-palladium before SEM examination.3.

The thicknesses of the films were measured from SEM micrographs of the cross sections and verified by a caliper. The samples were coated with gold-palladium before SEM examination.3. Results and discussionEffect of Cu additiveFigure 2 shows the surface morphologies of the pure PC films sintered at different scanning speeds. For BS = 100%, figure 2a, many voids are present due to incomplete fusion of the PC particles. As BS was reduced to 80 & 70%, i.e. increasing the laser energy density, the number of voids dropped and the surface became smoother. When BS was further reduced to 60%, however, there were some blemishes on the surface.

increasing the laser energy density, the number of voids dropped and the surface became smoother. When BS was further reduced to 60%, however, there were some blemishes on the surface. These were signs of material degradation under an excessively high energy density of the laser beam.112a) 50% P, 100% BSb) 50% P, 80% BSc) 50% P, 70% BSd) 50% P, 60% BSFigure 2 SEM micrographs showing the surface views of PC films sintered at 50% of P anddifferent values of BS.Figure 3 shows the surface morphologies of the PC/Cu films. A higher Cu content led to a more porous surface structure, as if the films had been sintered at a lower energy density of the laser beam. For blends of PC/Cu (1/1), figures 3b&d, finer Cu powder gave more voids than the coarser Cu powder did. Perhaps, the larger amount of smaller Cu particles had adversely affected the flow of molten polymer during the sintering process.

Perhaps, the larger amount of smaller Cu particles had adversely affected the flow of molten polymer during the sintering process. In comparison, blends of PC/Cu (10/1), figures 3a&c, showed a very similar surface morphology despite the difference in particle size of the Cu additive. This was probably due to the fact that the Cu content was low and the effect of particle size was not apparent.Besides, the thickness of the films with a high Cu content was smaller than that of the films with a low Cu content, figure 4. The phenomenon was probably due to increase in heat loss after addition of Cu. In order to prove that increasing the heat loss (particularly through the powder bed), would reduce the film thickness, sintering was carried out on powder beds of different thicknesses. When the thickness of the powder bed, supported on a stainless steel base, was reduced from 4.9mm to 1.6mm, there was a corresponding drop in thickness of the sintered PC films, figure 5.

When the thickness of the powder bed, supported on a stainless steel base, was reduced from 4.9mm to 1.6mm, there was a corresponding drop in thickness of the sintered PC films, figure 5. It is also noteworthy that when sintering at P/BS ratios below 1 (the lower the P/BS ratio, the lower the ED), reducing the powder thickness from 4.9 to 3 mm did not cause any major change in the thickness of the sintered films. However, when P/BS ratio was above 1,113the sintered films became significantly different in thickness. One possible explanation is that at a low P/BS ratio, i.e. the rate of energy input is relatively low, a 3mm thick powder bed is sufficient to provide good insulation against heat loss through the stainless steel base. Further increase in the powder bed thickness only improves the insulation properties marginally. For a high P/BS ratio, however, the temperature gradient through the powder bed thickness might change significantly when it was reduced from 4.9mm to 3mm.

Plots of thickness against P/BS ratio for sintered films with different Cu contents andpowder sizes.1.701.501.6mm PC powder bed1.30)m m ( s s e n k c i h T1.103mm PC powder bed0.904.9mm PC powder bed0.700.500.400.600.801.001.201.401.601.80P/BS ratioFigure 5. Plots of thickness against P/BS ratio for sintered films with different powder bedthicknesses.a) PCb) 1:1 PC/Cu composite with Cu powders < 45mmFigure 6. SEM micrographs of sintered films, sintering parameters: P = 50%, BS = 50%, DPI = 500 and PPI = 500, white spots in (a) are signs of material degradation and those in (b) are Cu particles.115Effect of SiO2 and QuartzFigure 7 shows the surface views of three different films, i.e. PC, PC/SiO2 and PC/quartz, sintered under the same condition. The additives seem to have an adverse effect on the fusion of the polymer. In particular, the SiO2 powder led to severe degradation of the polymer on the film surface.

The additives seem to have an adverse effect on the fusion of the polymer. In particular, the SiO2 powder led to severe degradation of the polymer on the film surface. Polymers degrade under exposure to high temperature, oxygen and ozone, ultraviolet light, moisture, radiation and chemical agents[5]. In this case, degradation is likely due to high temperature generated on the film surface. The chemical composition of SiO2 and quartz are basically the same and their thermal properties are extremely similar. The different morphologies of the two films, figures 7b&c, were probably caused by the different amounts of energy dissipated on the surface during the sintering process. According to Karapatis[6], with a suitable size ratio between coarse and fine particles (over 10:1) and adequate composition (30%) of fine particles, the powder bed density can increase by 15%.

According to Karapatis[6], with a suitable size ratio between coarse and fine particles (over 10:1) and adequate composition (30%) of fine particles, the powder bed density can increase by 15%. Adding fine SiO2 powder will partially fill up the voids of the powder bed, hence reducing the penetration depth of the laser beam, and more energy was dissipated near the surface. Table 2 shows a decrease in porosity after adding additives. More importantly, the fine SiO2 powder (32mm) had a larger total surface area than the coarse quartz powder (140mm). This would enhance interaction with the laser beam.a) PCb) 20% SiO2c) 20% QuartzFigure 7.

This would enhance interaction with the laser beam.a) PCb) 20% SiO2c) 20% QuartzFigure 7. Effect of quartz and SiO2 on surface morphology of sintered PC films, sinteringparameters: P = 30%, BS = 50%, DPI = 500 and PPI = 500.116Table 2 Bulk density and porosity of PC, PC/SiO2 and PC/quartz.Materials Bulk density (g/cm3) Volume occupied by PC (%) Volume occupied by additive (%) Porosity (%)PC 3.77 60.8139.19PC/SiO2 4.602 59.38 3.60 37.02PC/quartz 4.984 64.31 3.78 31.91Some finer quartz powder (<45mm) was used to blend with PC and sintered under similar conditions. The result is shown in figure 8. It is now obvious that the finer additives indeed caused heavier degradation of the polymer near the surface of the sintered films. A thickness measurement, figure 9, shows that the PC/SiO2 and PC/quartz films are thinner than the pure PC film.

A thickness measurement, figure 9, shows that the PC/SiO2 and PC/quartz films are thinner than the pure PC film. The phenomenon is in agreement with the argument that most of the laser energy was dissipated near the surface of the composite films.a) 20% SiO2b) 20% QuartzFigure 8. Effect of quartz (<45mm) and SiO2 (32 mm) on surface morphology of sintered PCfilms, sintering parameters: P = 35%, BS = 50%, DPI = 500 and PPI = 500.1.000.90)m m0.80(s s e n k c i h T0.700.60PCSiO2 Quartz0.500.400.400.500.600.700.800.901.00P/BS ratioFigure 9 Thickness change of sintered films with SiO2 and quartz.1174. ConclusionsAll three additives had an adverse effect on the fusion behavior of PC and resulted in a more porous surface structure of the sintered films. This can be attributed to increase in viscosity of the polymer melt because the additives remained solid during the laser sintering process.

This can be attributed to increase in viscosity of the polymer melt because the additives remained solid during the laser sintering process. For a given additive content, a smaller particle size would increase energy dissipation near the surface of the powder bed because of the enhanced area of contact with the laser beam. However a smaller particle size of the additives would hinder fusion of molten polymer more significantly, probably due to the increased area of the polymer/additive interface. The Cu powder has a considerably higher thermal conductivity than the PC powder, which helps to improve the overall heat transfer properties and hence alleviates the phenomenon of polymer degradation on the film surface. However, the increased rate of heat loss resulted in a reduction in the thickness of the sintered films. In contrast, SiO2 and quartz caused degradation of the polymer.

However, the increased rate of heat loss resulted in a reduction in the thickness of the sintered films. In contrast, SiO2 and quartz caused degradation of the polymer. The phenomenon was likely caused by the increased energy dissipation near the film surface and poor heat transfer properties of these two additives.References1. Badrinarayan, B. and Barlow, J.W., Effect of Processing Parameters In SLS of Metal- Polymer Powers, Solid Freeform Symposium Proceedings, 1995, pp. 55-63.2. Incropera, F.P. and DeWitt, D.P., Fundamentals of Heat and Mass Transfer, John Wiley & Sons: New York, 2nd edition, 1985, pp. 740-782.3. Edenbaum, Jesse, Plastics Additives and Modifiers Handbook, Van Nostrand Reinhold, 1992, pp. 507, 525, 533.4. Nelson, J.C., Selective Laser Sintering: A Definition of the Process and an Empirical Sintering Model, Ph.D. Dissertation, University of Texas at Austin, 1993.5. Fried, Joel R., Polymer Science and Technology, Prentice Hall, 1995, pp.

fects of Degree of Particle Melt and crystallinity in SLS Nylon-12 partsC.E.Majewski*, H.Zarringhalam† and N.Hopkinson* *Rapid Manufacturing Research Group, Loughborough University, Loughborough, United Kingdom. LE11 3TU, † Materials Solutions, University of Birmingham Campus, Edgbaston, Birmingham, United Kingdom. B15 2TTReviewed, accepted September 10, 2008 AbstractDifferential Scanning Calorimetry (DSC) traces for SLS Nylon-12 parts display two distinct melt peaks, which have been related to the presence of both melted and crystallised regions, and un-melted particle cores within the part. The relative proportions of each region are defined by the term ‘Degree of Particle Melt’ (DPM), and have a large effect on the mechanical properties of a part. This paper demonstrates that the % crystallinity of SLS Nylon-12 parts is dependent on the DPM.

This paper demonstrates that the % crystallinity of SLS Nylon-12 parts is dependent on the DPM. Crucially, research has also shown that the trends for some tensile properties (notably Tensile Strength and Young’s Modulus) change once full melting is complete.1.0IntroductionPrevious research1,2 identified that, for Selective Laser Sintering (SLS) of polymers, only partial melting occurs, whereby larger particles receive insufficient energy to completely melt.

Figure 1 clearly illustrates the presence of particle cores (un-melted material), surrounded by melted areas.Figure 1 - Microstructure of SLS material31.1Degree of Particle MeltDifferential Scanning Calorimetry (DSC) traces for laser sintered Nylon-12 have been shown to include two distinct melt peaks, which have been related to the melted and un-melted regions within a part (see Figure 2).0 0) ) g g / / W W( (w w o o l l f ft t a a e e H H1 -12 -23 -3Virgin powder Virgin powder Part Part4 -45 -5Temperature (°C) Temperature (°C)160 160170 170180 180190 190200 200Temperature (°C) Temperature (°C)181°C 188°C 189°C 181°C 188°C 189°CUn-melted particle core Un-melted particle coreMelted & crystallised region Melted & crystallised regionFigure 2 Relationship between DSC peaks and material microstructure (adapted from3)It has also been shown previously that the percentage crystallinity of an SLS part is dependent on the DPM of that part.

It is known that, for traditionally processed semi-crystalline materials, an increase in the crystalline content leads to an increase in tensile strength and stiffness, but at the detriment of the ductility, as it is the amorphous regions of a polymer that provide it’s ability to yield without breaking7. Given the correlation identified between DPM and mechanical properties, it was considered logical to expect that the DPM would have an effect on the percentage crystallinity of SLS parts.2.0 Experimental ProcedureTo establish the effect of DPM and crystallinity on mechanical properties, a series of parts was produced under different build conditions, in order to provide a range of differing DPM, as shown in previous research6.DSC was performed for each build, and the crystallinity calculated as described in Section 2.2.1.

Tensile testing was performed from each set of build conditions, and the measured properties were compared with the crystallinity.2.1SLS Part ProductionIn order to produce parts with varying levels of DPM for this research, parts were produced on a single SLS machine (3D Systems SLS Vanguard), and parameters which would specifically alterthe energy input were varied. The material used was Duraform PA (Nylon-12). Specimens with dimensions of 180 x 10 x 4 mm were produced, from which DSC measurements were performed.Six tensile specimens were produced in each build according to ISO-527-2:1996 type 1A, as shown in Figure 3. Parts were built in the x-axis to reduce end of vector effects, and built flat to reduce build time. The layer thickness used was the standard 0.1 mm.

Parts were built in the x-axis to reduce end of vector effects, and built flat to reduce build time. The layer thickness used was the standard 0.1 mm. In addition to the tensile specimens, one ‘sample’ part was built to obtain specimens for DSC testing as required.z zy yx x4mm 4mm80mm 80mm10mm 10mm200mm 200mmFigure 3 - ISO 527-2:1996 type 1A tensile specimen parts and ‘sample’ partStandard build parameters for Duraform PA were used as default values, and for subsequent builds a single parameter was changed from the default, either to a high or a low level.

Table 1 shows the main build parameters which were varied, where Green indicates a higher level than default, and Red a lower value.Build1 – Default 2 – PHSP high 3 – PHSP low 4 – SFSS high 5 – SFSS low 6 – FLP high 7 – FLP low 8 – FSC high (2) 9 – FSC high (3)Part heater set point (PHSP)148 150 146 148 148 148 148 148 148Slicer fill scan (SFSS) spacing (mm) 0.15 0.15 0.15 0.13 0.17 0.15 0.15 0.15 0.15Fill power (FLP) (W) 11.5 11.5 11.5 11.5 11.5 13.5 9.5 11.5 11.5laserFill count (FSC) 1 1 1 1 1 1 1 2 3Table 1 - Machine parametersscan2.2 Material Testing2.2.1 Differential Scanning Calorimetry (DSC)DSC measurements were taken using Shimadzu DSC 60WS equipment, and analysed using TA60 software, also from Shimadzu. Sample masses were 5 mg +/- 0.05 mg, and samples were heated from ~20 °C to 220 °C at a rate of 10 °C per minute.

Sample masses were 5 mg +/- 0.05 mg, and samples were heated from ~20 °C to 220 °C at a rate of 10 °C per minute. Samples were held at 220 °C for two minutes, and then cooled at 10 °C per minute.Figure 4 shows an example DSC graph, with the relevant area of interest highlighted, andFigure 5 indicates the exact positions from which measurements were taken.See Figure 5 See Figure 5Figure 4 - Example DSC traceFigure 5 - Position of measurementsThe percentage crystallinity was determined by dividing the heat of melting (indicated in Figure 5) by the known heat of melting for a 100% crystalline specimen. This value was taken to be 209.3J/g, as reported by Gogolewski et al8.2.2.2 Tensile TestingIn order to test the tensile properties, a Zwick testing machine with long travel contact extensometer was used.

E-Modulus was measured using a 1 mm/min strain rate, and Tensile Strength and Elongation at Break were measured at 5 mm/min.3.0 Results3.1Percentage CrystallinityFigure 6 shows the % crystallinity calculated for each build.Figure 6 - % Crystallinity for each buildIt can be seen that each high value of a parameter provided lower crystallinity than the default, indicating a higher DPM, and vice versa. Previously it has been shown that the point at which melting becomes complete occurs between 25.8 and 24.9 % crystallinity, falling between the lowest single scan point and the double scan4. From this point forward, parts containing only fully melted material will be referred to as ‘single phase’, and those containing both melted and un-melted material will be referred to as ‘double phase’.3.2Tensile PropertiesThe following figures show the Elongation at Break (Figure 7) and the Tensile Strength (Figure 8) plotted against the % crystallinity.

Young’s Modulus has previously been shown6 as being affected in no significant manner by the DPM, and therefore has not been analysed here.It should be noted that the blue dashed line does not represent the actual boundary between the double phase and single phase regions. The actual boundary could lie anywhere between the last data point in the double phase region and the first data point in the single phase region. Therefore for each region the trend line is extrapolated into the neighbouring region. These charts show a striking difference in mechanical properties between the double phase material (cores) and single phase material (no cores).3.2.1 Elongation at BreakFigure 7 - Effect of % crystallinity on EaBIt can be seen from Figure 7 that, as the % crystallinity decreases, the Elongation at Break increases, as would be expected from the discussions presented in Section 1.1.

However, once DPM is complete, there appears to be a definite step change in the values recorded, indicating that the effect of crystallinity changes once full melting is achieved.3.2.2 Tensile StrengthFigure 8 - Effect of crystallinity on Tensile StrengthFigure 8 shows that, as % crystallinity decreases, the Tensile Strength increases, until the point at which DPM is complete. At this point, there appears to be an instantaneous ‘tipping point’, at which the direction of the trend is reversed.3.2.3 SummaryIt can be concluded from these results that the % crystallinity has an effect on the tensile properties, during the period in which melting of particles is not complete. Once melting is fully complete, a new trend can be observed.

Once melting is fully complete, a new trend can be observed. In effect this leads to a distinction between fully melted and partially melted SLS parts, indicating that it may be appropriate to treat them as being two different materials – a homogenous, single phase material, and a quasi-composite material comprised of both melted and un-melted Nylon-12.Whilst further work would be required to explain the apparent changes in the mechanical property results once melting is complete, it is possible that this is due to differences in the rate of crystallisation caused by the presence or absence of the un-melted particle cores, and therefore corresponding differences in the size of spherulites within the processed material.4.0 Use of % crystallinity to quantify DPMThe % crystallinity of an SLS Nylon-12 part will be the sum of the % crystallinity of the un-melted cores and the melted and re-crystallised material. Figure 9 shows this in diagrammatic form.

Figure 9 shows this in diagrammatic form. The y axes indicate the relative percentages of the part which are comprised of melted and crystallised material (MCM) and un-melted, core material. The x axes correspond to the crystallinity of each phase. The % crystallinity of the un-melted material is assumed to be that ofthe powder itself, previously calculated to be 47 %. The % crystallinity of the fully melted material is taken to be 25 %, as determined in previous research4.

Four different example compositions are demonstrated, and schematic illustrations indicating their approximate composition are included.Figure 9 - Total % crystallinity as the sum of Core and MCM % crystallinityThe total % crystallinity can therefore be calculated using Equation 1 below.Total Crystallinity = (%MCM x MCM crystallinity) + (%Core x Core crystallinity)Equation 1Taking % core to be 1 - % MCM gives:Total Crystallinity = (%MCM x MCM crystallinity) + ((1-%MCM) x Core crystallinity)Total Crystallinity = (%MCM x MCM crystallinity) + Core crystallinity - (%MCM x Core crystallinity)Total Crystallinity – Core crystallinity = %MCM x (MCM crystallinity - Core crystallinity)DPM (% MCM) = (Total crystallinity – Core crystallinity) / (MCM crystallinity – Core crystallinity)Equation 2Assuming knowledge of the % crystallinity of the base powder, and that of fully melted material, it therefore follows that it is possible to calculate the DPM of the part using its crystallinity as measured from a DSC chart.

The small size of specimen required for performing DSC would mean that it is entirely practical to include a sacrificial specimen in every build, in order to determine repeatability of melting, and therefore of mechanical properties. This is a far more practical approach than the destructive testing of parts, or the inclusion of much larger test specimens throughout builds. Further investigations, possibly involving varying only a single parameter (e.g.

Further investigations, possibly involving varying only a single parameter (e.g. laser power) could be carried out in order to confirm the accuracy and repeatability of this method.5.0 ConclusionsThis work has shown that the percentage crystallinity of a two-phase SLS part, as calculated from a DSC chart, has an appreciable effect on the Tensile Strength and Elongation at Break, whereby a decrease in crystallinity leads to an increase in the mechanical property.Crucially it has been shown that, once melting is complete, the trends become different, indicating that the material the properties and dependencies can be treated as though for an entirely new material.

Further investigation in this area will be required in order to establish significance.It has also been hypothesised that, given knowledge of the percentage crystallinity of the fully melted material, and the base powder, the relative percentages of the material within a part which are fully melted can be calculated by determining the % crystallinity of the part from a DSC trace.Further work could focus on the use of imaging software to correlate the quantified DPMvalues with optical microscopy images in order to further verify this hypothesis.6.0 AcknowledgementsThis research was carried out by Hadi Zarringhalam with support from the PowderMatrixFaraday Partnership, Solid Concepts Inc. and 3D Systems UK7.0 References1 Moeskops, E., Kamperman, N., van de Vorst, B.

Naval Air Systems Command, Patuxent River, MD 20670Corresponding author: shamsaei@auburn.eduAbstractIn this study, the effects of part size on thermal history and mechanical properties of additively manufactured 17-4 PH stainless steel were investigated under monotonic tensile and strain-controlled fatigue loadings. Two sets of specimens were machined from square rods and oversized specimens, which were fabricated using a laser bed powder fusion (L-PBF) process, to introduce variation in specimen geometry and consequently thermal history. Monotonic tensile tests were conducted at a strain rate of 0.001 s-1. Fully-reversed (Rε = -1), strain-controlled fatigue tests were performed at 0.003 and 0.0035 mm/mm, and varying test frequency to maintain a constant average strain rate in all tests. Experimental results indicated minimal effect of specimen geometry on the tensile properties of L-PBF 17-4 PH SS, which were also found to be comparable to the wrought material.

Experimental results indicated minimal effect of specimen geometry on the tensile properties of L-PBF 17-4 PH SS, which were also found to be comparable to the wrought material. On the other hand, some influence of specimen geometry on fatigue behavior was observed. Specimens machined from square rods exhibited slightly higher fatigue resistance as compared to specimens machined from oversized specimens. Furthermore, thermal simulations demonstrated higher bulk heating in specimens machined from oversized specimens as compared to those from square rods, which indicated the effect of part geometry on thermal history experienced by the fabricated part.Keywords: Fatigue; Thermal Simulation; Additive manufacturing; Laser-based powder bed fusion (L-PBF)IntroductionAdditive manufacturing (AM) technologies have enabled designers to fabricate application-specific, customized parts with complex geometries.

With the capability of producing complete assemblies, AM is also being considered as a cost-effective manufacturing process as it minimizes the number of part assemblies and cost. Due to these advantages, AM has transitioned from a traditional means of rapid prototyping to a tool for advanced manufacturing [1]. There is1277also an increased interest in recent years from academia, private industries, and governmental sectors towards AM technologies in effort to fabricate high quality parts with optimal mechanical properties.Although AM possesses various advantages, the susceptibility of AM parts to inherent defects, such as pores resulting from entrapped gases and voids from lack of fusion (LoF) between the subsequent layers, is one of the major drawbacks of AM processes.

As the fatigue resistance of structural parts/components are sensitive to defects, the trustworthiness of AM parts is not fully established, particularly for load bearing and fracture-critical parts in aerospace and automotive applications. Furthermore, it is known that thermal history experienced by AM parts during the manufacturing process is directly affected by the combination of process parameters, such as laser power, hatch spacing, layer thickness, scan speed, etc. Each combination of these parameters can result in variation of parts’ thermal history, which, in turn, directly affects their microstructural characteristics, and ultimately mechanical properties of parts [2-4].In addition, size and geometry of fabricated parts can influence thermal history. Part size can directly dictate the amount of solidified area available for the heat to dissipate through the previously deposited layers.

Part size can directly dictate the amount of solidified area available for the heat to dissipate through the previously deposited layers. The geometry of the fabricated part can influence the amount of heat dissipating though previously solidified layers and the powder bed, which could result in different thermal histories even within the same layer. Therefore, a combination of different process parameters along with part size and geometry can locally affect cooling rate, temperature gradient, as well as bulk heating during the fabrication. As a result, a thorough study on the effect of part size and orientation is necessary to accurately relate the mechanical behavior of test specimens to a large component with complex geometries, with the end goal of establishing process-structure- property-performance relationship between the test specimens and actual components [5].Weingarten et al.

[6] studied the effects of various process parameters on hydrogen- entrapped pore density in Al-Si-10Mg alloy, which was fabricated using a laser powder bed fusion (L-PBF) process. Among various process parameters including scan speed, hatch spacing, and laser beam diameter, which governs the time between melting and solidification (tM) of melt pool was reported to have the most significant effect on pore density. As tM is inversely related to the scan speed, higher scan speed results in lower tM. This decreases the time for hydrogen (present as moisture in the powder) to escape from the melt pool, and increases the concentration of hydrogen that transforms into pores during the solidification process.Similarly, Yang et al. [7] also conducted a study to obtain the effects of process parameters and part geometry on the mechanism of pore formation in L-PBF Al-7Si-Mg and L-PBF Al-10Si- Mg. The volumetric (i.e. energy per unit volume) and areal (i.e.

The volumetric (i.e. energy per unit volume) and areal (i.e. energy per unit area) energy density, which are functions of laser power, scan speed, hatch distance, and layer thickness were related to the formation of different types of defects including, LoF defects, small gas pores, and large keyhole-type pores. Larger LoF defects were attributed to the lower volumetric energy density resulting from higher scan speed. On the other hand, formation of small gas pores were seen to be affected from moisture present in the powder and high areal energy density. Large round pores were reported to be associated with the keyhole mode of melting due to high volumetric energy density. Furthermore, the effect of part geometry on the distribution of defects was investigated by fabricating a special type of specimen that consists of cubic geometry at the bottom, cylindrical in the middle section, and thin plate at the top.

It was found that the thin plate section contained higher amount of porosity level as compared to the other two sections. This was1278attributed to the smaller cross-section area available for heat dissipation as well as use of meander scan speed [7]. Hengfeng et al. [8] investigated the effect of energy density as a function of laser power, hatch spacing, and layer thickness on the porosity of AM PH 17-4 SS. It was found that, at high scan speed and laser power, pores were formed due to balling phenomenon and thermal stress cracking, while low laser power and scan speed produced LoF crevices due to insufficient melting [8].Recently, Pegues et al. [9] studied the effect of specimen surface area and gage section radius on fatigue behavior of diagonally-built AM Ti6Al4V alloy in as-built condition. Results showed that the gage section radius had more effect on fatigue resistance as compared to the surface area.

Results showed that the gage section radius had more effect on fatigue resistance as compared to the surface area. Furthermore, as specimens were fabricated at 45° with respect to the substrate, the majority of heat is dissipated from the previously solidified layer in the top-skin surface (surface facing away from the substrate), while for the down-skin (surface facing towards the substrate) surface, heat is mostly dissipated through powder. As the heat transfer coefficient for solidified part is higher than the powder, there is a distinction of the thermal history between the top- and down-skin surfaces, leading to difference in surface roughness.

Additionally, fracture surfaces also revealed that cracks always initiated from the down-skin surface that had higher surface roughness, indicating a significant effect of part geometry on the corresponding thermal history variation, and consequently fatigue behavior.To better understand the fatigue behavior in AM materials, the formation of defects resulting from thermal history experienced by the fabricated parts should be investigated. In this study, the effects of part geometry on monotonic tensile and fatigue behavior of L-PBF 17-4 precipitation hardening (PH) stainless steel (SS) are investigated in attempt to relate the variation and distribution of defects. The material in this study, 17-4 PH SS, is currently of interest in various applications due to its corrosion resistance coupled with high yield and ultimate tensile strengths. These include aerospace (particularly in turbine blade designs), marine, nuclear power plants, as well as in oil and gas industries.

These include aerospace (particularly in turbine blade designs), marine, nuclear power plants, as well as in oil and gas industries. Following the introduction in this study, the material and experimental procedures are presented. The experimental results are then discussed along with the fractography analysis and thermal simulations. Finally, important conclusions are drawn.Material and Experimental ProceduresArgon atomized 17-4 PH SS powder produced by LPW Inc. with size distribution ranging from 15-45 µm was used to fabricate specimens using an EOS M290, a L-PBF system. The employed process parameters are shown in Table 1. Furthermore, shape distribution of the utilized powder was also investigated using a scanning electron microscope (SEM) and shown in Fig. 1(a).

The employed process parameters are shown in Table 1. Furthermore, shape distribution of the utilized powder was also investigated using a scanning electron microscope (SEM) and shown in Fig. 1(a). SEM image revealed that the majority of the powder particle were spherical in shape with average size of 25 µm.Table 1: Process parameter used to fabricate the L-PBF PH 17-4 SS specimensLaser power (W)Scanning speed (mm/s)Hatching distance (μm)Layer thickness (μm)220755.5100401279The two sets of specimens, one from vertically-built L-PBF PH 17-4 SS square rods with dimensions 12x12x90 mm shown in Fig. 1(b) and another set of vertically-built L-PBF PH 17-4 SS oversized specimens with gage diameter of 8 mm in Fig. 1(c), were machined using a CNC lathe. The specimen dimensions with uniform gage section, following ASTM E606 [10] standard, are presented in Fig. 2.

1(c), were machined using a CNC lathe. The specimen dimensions with uniform gage section, following ASTM E606 [10] standard, are presented in Fig. 2. For simplicity, the specimens fabricated from square rods and oversized specimens will be referred to as Specimen Set A and Specimen Set B, respectively, throughout this manuscript. All specimens were subjected to CA H 1025 heat treatment, which was selected based on a study of effects of different heat treatment on the mechanical behavior of 17-4 PH SS, which is currently under preparation.

To eliminate the effect of surface roughness, specimens were further polished using different grit sandpapers to remove all the horizontal machining marks.1280Figure 1: (a) SEM image of PH 17-4 SS powder along with L-PBF PH 17-4 SS parts built in vertical direction with two different specimen geometries including (b) 12x12x90 mm square rod and (c) oversized specimen with gage diameter of 8 mm.Figure 2: Specimen geometry for monotonic tensile and fatigue tests fabricated based on ASTM E606 standard [10].The mechanical testing, including monotonic tensile and fatigue tests, were performed using a servo hydraulic MTS landmark test frame with the load capacity of 100 kN. All tests were conducted in strain-controlled mode, and the axial strain was measured and controlled using a mechanical extensometer with 10 mm gage length.

All tests were conducted in strain-controlled mode, and the axial strain was measured and controlled using a mechanical extensometer with 10 mm gage length. As elongation to failure of 17-4 PH SS exceeded the allowable travel distance of the extensometer, monotonic tensile tests were first conducted at strain rate of 0.001 s-1 under strain-controlled mode, and later switched to displacement-controlled mode. Fully-reversed (Rε = -1) uniaxial strain-controlled fatigue tests were carried out at 0.003 and 0.0035 mm/mm strain amplitudes. The test frequencies were varied to maintain similar strain rate in all fatigue tests to eliminate any strain rate effect.Physical Model and Numerical MethodA major challenge in simulating the AM process is to develop an efficient model that can produce information and knowledge at the overall scale of the desired part and build-process to inform engineering decisions [11, 12].

There can be kilometers of laser travel length and thousands of powder layers in a single build, thus presenting a computational challenge. To be useful, part- scale simulations must be executable in a reasonable amount of time, while still retaining sufficient physical fidelity to yield trustworthy results. An eventual goal is to have efficient process models that could be included as a part of an automated process optimization.At the part scale, one neglects the details of local laser-powder interaction. Instead, the goal is to describe part thermomechanical behavior both during and after the build is completed. In this approximation, the powder was represented as a reduced-density, low-strength solid. The deposition of the laser energy into the powder was then represented by a volumetric energy source term such as that derived by Gusarov et al. [13]. Melting was represented thermally through a latent heat parameter and mechanically as a near-total loss of strength.

[13]. Melting was represented thermally through a latent heat parameter and mechanically as a near-total loss of strength. The “mushy zone” at the1281melt pool boundary was represented by having its temperature-dependent strength rise as temperature falls below the solidus temperature.For modeling, each layer is typically simulated separately with the per-island scan strategy. In the proposed numerical method, the heat flux at the top layer was assumed to be uniform. The total amount of time that the heat flux is applied was based on the time required to build one layer. In the case of rectangular section shown in Fig. 3(a), the time period was calculated as:𝑚𝑚 × 𝑙𝑙 𝑤𝑤 × 𝑣𝑣 where l is the length of the rectangle, m is the thickness of the part, w is the hatching space, and v is scanning speed. Similarly, the total time to fabricate a slice of a cylindrical part with circular cross-section of diameter of D as shown in Fig.

Similarly, the total time to fabricate a slice of a cylindrical part with circular cross-section of diameter of D as shown in Fig. 3(b), is calculated as:𝑡𝑡 =22𝑡𝑡 =𝑚𝑚 × 𝑙𝑙 𝑤𝑤 × 𝑣𝑣𝜋𝜋 𝐷𝐷 4� 𝑚𝑚 × 𝑙𝑙=𝜋𝜋 𝐷𝐷 4� 𝑤𝑤 × 𝑣𝑣Figure 3. Schematic of built area for (a) the square rod and (b) the oversized specimen.The constant heat flux applied at the top of the layer,, is calculated as:𝒒𝒒pb,i𝒒𝒒pb =𝑃𝑃 𝑚𝑚 × 𝑙𝑙𝑃𝑃 2 𝜋𝜋 𝐷𝐷 where P is laser power. Another assumption used to facilitate the simulation was to attach layers together to form a bulk-layer, which significantly decreased the number of layers required to model the AM part.𝒒𝒒pb =4�1282(1)(2)(3)(4)Experimental Results and DiscussionsEngineering stress-strain responses obtained from the monotonic tensile tests for two sets of L-PBF PH 17-4 SS specimens are displayed in Fig. 4.

4. As explained in the previous section, the stress-strain responses were captured up to the extensometer’s limit prior to switching from strain- to displacement-controlled mode. The corresponding tensile properties are shown in Table 2 and compared to those of wrought material reported in [14]. As seen in Fig. 4, the effect of specimen geometry on the tensile behavior is not significant.

As seen in Fig. 4, the effect of specimen geometry on the tensile behavior is not significant. Tensile properties for L-PBF PH 17-4 SS specimens were comparable to the wrought material (10% difference in yield strength and 16% difference in ultimate tensile strength).Figure 4: Engineering stress-strain responses for L-PBF PH 17-4 SS specimens.Table 2: Tensile properties of L-PBF and wrought PH 17-4 SS specimens.Specimen Geometryσy (MPa)σu (MPa)E (GPa)L-PBF Specimen Set A11701203205L-PBF Specimen Set B11601188202Wrought [14]10701000200To obtain the failure mechanism and defect distribution in specimens, fractography analysis was conducted on the failed tensile specimens. Fracture surface for Specimen Set A and Specimen Set B are shown in Figs. 5(a) and 5(b), respectively. Both types of specimen geometry1283exhibited a cup and cone type fracture surface as well as dimples, indicating a ductile failure.

5(a) and 5(b), respectively. Both types of specimen geometry1283exhibited a cup and cone type fracture surface as well as dimples, indicating a ductile failure. Moreover, these dimples can also show the presence of precipitates common in PH SS. The Cu- rich precipitates become coarser and non-coherent to the matrix due to high temperature heat treatment for longer period of time (i.e. CA-H1025). As a result, dislocations must bypass these coarse non-coherent precipitates, which can lead to difference in the deformation rate of matrix and precipitates and consequently formation of dimples [15]. Additionally, some facets representing brittle fracture behavior as pointed out by the white arrows in Fig. 5 were also seen on the fracture surfaces. These observations may suggest that the failure mode of L-PBF 17-4 PH SS under monotonic tensile loading may be a combination of both ductile and brittle fracture. Furthermore, some secondary cracks, denoted by the red solid arrows in Fig.

Furthermore, some secondary cracks, denoted by the red solid arrows in Fig. 5(a), were observed only in Specimen Set A, while they were absent in Specimen Set B. However, these cracks seem to have minimum effect, as the tensile properties as both sets of specimen were similar, as seen in Fig. 4, and comparable to the wrought counterpart.(a)(b)Figure 5: Fracture surfaces for specimens fabricated from (a) Specimen Set A and (b) Specimen Set B failed under monotonic tensile loading.Although the presence of defects may not have a substantial impact on the monotonic tensile properties, defects can significantly affect the fatigue behavior of AM materials, which is more defect-driven. Such observation was also noted in the previous study on L-PBF 316L SS [16]. Fatigue lives for two sets of specimens under uniaxial, fully-reversed strain-controlled1284condition are displayed in Fig 6.

Fatigue lives for two sets of specimens under uniaxial, fully-reversed strain-controlled1284condition are displayed in Fig 6. As seen, Specimen Set A exhibited better fatigue resistance when compared to Specimen Set B, approximately by a factor of 3 and 1.5 in lower (0.003 mm/mm) and higher (0.0035 mm/mm) strain amplitude tests, respectively. More scatter in fatigue data at lower strain amplitude is also evident in this figure. Such result can be attributed to the fact that, at lower strain amplitudes the failure mechanism is more defect driven compared to the high strain amplitude loading [17]. Therefore, fatigue lives at lower strain amplitude are more sensitive to location, size, and types of defects, which can add scatter in the fatigue data.

Therefore, fatigue lives at lower strain amplitude are more sensitive to location, size, and types of defects, which can add scatter in the fatigue data. Nonetheless, a distinct variation in the fatigue lives between Specimen Set A and Set B can be observed in Fig 6, suggesting the effect of part geometry on fatigue life of L-PBF 17-4 PH SS specimens.Figure 6: Strain-life data showing variation in fatigue life of 17-4 PH SS specimens.Figures 7(a) and 7(b) display the fracture surface of Specimen Set A (317,060 reversals) and Specimen Set B (145,330 reversals), respectively, both subjected to strain amplitude of 0.003 mm/mm. As can be seen, the fatigue crack in Specimen Set B initiated from two gas pores at the specimen’s surface. On the other hand, for the Specimen Set A, crack was observed to initiate from a single pore, also at the surface of the specimen.

On the other hand, for the Specimen Set A, crack was observed to initiate from a single pore, also at the surface of the specimen. Similar observation with the crack initiating from multiple pores was also seen in another specimen from Specimen Set B, which may have resulted in lower fatigue resistance as compared to Specimen Set A. It is to be noted that X-ray computed tomography should be conducted to determine the distribution of defects in each set of specimens, followed by further fractography analysis to accurately conclude the results.1285(a)(b)Figure 7: Fracture surfaces showing crack initiating factors for specimens fabricated from (a) oversize specimen (Specimen Set A) and (b) square rod (Specimen Set B) subjected to 0.003 mm/mm strain amplitude.Thermal Simulation and DiscussionsAs mentioned previously, the variation in thermal history experienced by the fabricated parts can affect the distribution of defects.

Therefore, using a thermal simulation tool, the temperature distribution and heat accumulation for two types of specimen geometries was also investigated in this study. Temperature distribution after fabrication of oversized specimens (i.e. Fig 1(c)) and square rods (i.e. Fig. 1(b)), are presented in Fig. 8. It is observed that the temperature at newly deposited layers are higher compared to previously deposited layers. Additionally, higher accumulated heat was seen in the case of the oversized specimen compared to the square rod.1286Figure 8. Temperature distribution of fabricated (a) oversized specimen and (b) square rod after fabrication of the parts.Distribution of predicted temperature at different heights of the fabricated parts is shown in Figs. 9(a) and 9(b), respectively, for oversized specimen (i.e. Fig (1(c)) and square rod (i.e. Fig. 1(b)). As can be seen, after the fabrication, the maximum temperature for oversized specimen was C higher compared to the square rod.

Fig (1(c)) and square rod (i.e. Fig. 1(b)). As can be seen, after the fabrication, the maximum temperature for oversized specimen was C higher compared to the square rod. Such a variation in bulk temperature of the almost 100 fabricated part can affect the temperature gradient and cooling rate, which can have a significant effect on microstructural features and defect distribution. Hence, variation in the thermal history observed in Fig. 8 may be responsible for differences in fatigue lives exhibited by Specimen Set A and Set B in Fig. 6.⁰Figure 9. Temperature distribution of (a) oversized specimen and (b) square rod along the main axis after fabrication of last layer.1287ConclusionsIn this study, effects of part geometry on the tensile and fatigue properties of L-PBF 17-4 PH SS were investigated for specimens machined from parts with different sizes. Based on the experimental results and simulations, the following conclusions were drawn:1.

Based on the experimental results and simulations, the following conclusions were drawn:1. L-PBF 17-4 PH SS specimens machined from square rods and oversized specimens exhibited similar tensile properties, which were also comparable to those reported for the wrought material.2. Specimens machined from square rods generally possessed better fatigue resistance when compared to the specimens machined from oversized specimens.3. Thermal simulation results indicated difference in heat accumulation between the oversized specimens and square rods. At the end of the fabrication, the oversized specimens were shown to be approximately 100 °C higher in bulk temperature compared to the square rods.

At the end of the fabrication, the oversized specimens were shown to be approximately 100 °C higher in bulk temperature compared to the square rods. Such variation in thermal history can have significant effects on microstructure, defect distribution, and ultimately fatigue life of the L-PBF 17-4 PH SS.Results from the current study show that the fractography analysis and thermal simulation alone cannot completely explain the effect of part geometry on the fatigue resistance of L-PBF 17- 4 PH SS. Therefore, the relationship between part geometry and thermal history should be further analyzed using microstructural studies to determine the difference in grain structure, and x-ray CT scans to determine the distribution of defects resulting from variations in thermal history.AcknowledgementThis material is based upon work supported by the Naval Air Systems Command (NAVAIR).References[1][2][3][4][5][6]E.C. Santos, M. Shiomi, K. Osakada, T.

At extreme conditions, NASA HR1 possess high strength, high fatigue resistance, and high resistance to corrosion and hydrogen embrittlement. The main applications include structural components and liquid rocket engine nozzles with internal cooling channels. NASA has produced HR1 using vacuum induction melting (VIM), a considerably expensive fabrication method. Aimed to explore other more affordable and accessible manufacturing methods, HR1 specimens were fabricated under different parameters using Laser-Powder Directed Energy Deposition (LP-DED) and were heat treated through stress relief, homogenization, solution treatment and aging. The feasibility of this AM process was investigated by evaluating mechanical and microstructural analysis on specimens.

The feasibility of this AM process was investigated by evaluating mechanical and microstructural analysis on specimens. This work finalizes with discussion and remarks on tensile and low-cycle fatigue properties and its relationship with microstructural features.IntroductionHR1 is mainly composed of Fe-Ni-Cr has been used in liquid engine rockets components such as cooling nozzles [1,2]. This alloy has been produced traditionally using vacuum induction melting [1]. Alternative manufacturing technologies are being explored to substitute this process, reducing costs. However, all manufacturing technologies provide mechanical properties that must be considered for its application to make sure this component will experience any deformation of fracture when being in service.

Laser Powder Direct Energy Deposition (LP-DED) is one of the additive manufacturing (AM) technologies being analyzed since it can reduce the cost mentioned before and it allows to create complex and large geometries. Before using DED as the default manufacturing technology for cooling nozzles, it is important to study the mechanical properties that HR1 acquires through this process and how this can be improved. The main purpose of this project is to evaluate the impact of using different laser power and angle deposition parameters on microstructure, tensile strength, and low cycle fatigue performance of NASA HR1 LP-DED. To317analyze these effects, NASA has created different plates deposited at two different power settings, including 1070 W and 2620 W but also the angle at which powder is being melted varied from 0, 20 and 30 degrees as represented in figure 1.

12 plates were printed in total by creating 2 plates of each power/angle combination, for instance, 1070 W 0 ̊, 1070 W 20 ̊, 2620 0 ̊, and so on. In addition, all these plates have been heat treated according to NASA HR 1 cycle which involved stress relief for 1.5 hours, homogenization for 6 hours in vacuum, solution treatment for 1 hour in vacuum and finally aging treatment.Figure 1 CAD model representation of plates printed at different angles.Materials and methodsMetallographyMetallographic preparation was required to see the microstructure of all different plates and compare their differences if found. All plates were sectioned in such a way planes XY (parallel to build lines) and XZ (build direction) could be observed. All samples were hot mounted, then, grinded from 240 to 600 grit to then be polished using diamond suspension of 9 µm, 3 µm, and 0.1 µm.

All samples were hot mounted, then, grinded from 240 to 600 grit to then be polished using diamond suspension of 9 µm, 3 µm, and 0.1 µm. Once a mirror surface was observed, these were taken to an optical microscope to appreciate the porosity differs between the ones deposited at 1070 W and 2620 W. Three images of each sample taken under the microscope were used to calculate their density using ImageJ software. Furthermore, etching was completed using Kallings #2 etchant for about three minutes to see the microstructure on both planes.Mechanical TestingAll plates have been sectioned as mentioned before to get different bars that were subjected to tensile and low cycle fatigue testing as shown in figure 2. These plates were sectioned and marked at W. M. Keck Center and then sent to Astro Laboratories in Los Angeles, CA to be tested there.

These plates were sectioned and marked at W. M. Keck Center and then sent to Astro Laboratories in Los Angeles, CA to be tested there. This laboratory oversaw machining the 3D printed plates until having four tensile bars and three fatigue bars of each power/angle combination with the appropriate dimensions and shape. It was reported that both mechanical testing was performed per ASTM E8-22 and ASTM e606-21 standards until fracture for tensile and low cycle fatigue respectively.318Results and discussionDefects and microstructure analysisPorosity was measured along XY plane in all specimens to compare the density percentage. It was observed that between those samples printed at 1070 W had similar densities ranging from 99.92% to 99.98%. However, those printed at 2620 W showed densities from 99.79% to 99.85% being slightly lower. As seen in figure 3, the number of pores (indicated with red arrows) and also, they looked bigger in diameter, approximately 5% bigger.

As seen in figure 3, the number of pores (indicated with red arrows) and also, they looked bigger in diameter, approximately 5% bigger. This may be an effect of gas getting entrapped easier in deeper melt pools when increasing wattage [2], which in this case, is more than double power.Figure 3 Left picture obtained from a 1070 W and right image gotten from 2620 W sample.As mentioned before, all plates (12 total, 2 of each power/angle combination) have gone through the same heat treatment. Therefore, it was expected that microstructure was similar in all plates. However, since the metal powder was deposited at different laser power, the solidification processes were different and thus the as-built microstructures [1]. As shown in literature, before heat treatment, a dendritic microstructure is appreciated being titanium rich in dark areas.

As shown in literature, before heat treatment, a dendritic microstructure is appreciated being titanium rich in dark areas. It is also shown that increasing power also increases the depth and width of melt pools since the cooling rate is lower allowing these melt pools to grow [3]. Heat treatment plays an essential role in this paper since it has been evaluated if even when all plates have been printed at different angles and power settings, they have gotten a similar microstructure after heat treatment and, therefore, mechanical properties are similar as well.

As shown in figure 4, all samples have an austenitic microstructure.319ABCDEFFigure 4 Microstructure of HR1 manufactured with LP-DED at different settings including A) 1070 W 0 Degrees, B) 1070 W 20 Degree, C) 1070 W 30 Degrees, D) 2620 W 0 Degree, E) 2620 W 20 Degree, F) 2620 W 30 DegreeIt is appreciated that dendrites have been completely removed in all samples by heat treatment and similar average grain size of 2.13 ASTM. Titanium might take longer to diffuse from grain boundaries to the bulk of the grains, making this diffusion dependent on the grain size [3]. In all cases, grain size ranged from 1.85 to 2.16 ASTM. In addition, annealed twinning (as indicated with red arrows) were found in all microstructures most likely due to the homogenization step. These twins contribute to increasing strength by pinning dislocations. Therefore, it is important to compare if they are present in all microstructures, which was the case.

These twins contribute to increasing strength by pinning dislocations. Therefore, it is important to compare if they are present in all microstructures, which was the case. Since all microstructures are similar in terms of grain shape and size but also the presence of annealed twinning, it was expected to see similar mechanical behavior.Mechanical TestingTensile behaviorDifferent average tensile properties, including 0.2% yield strength (YS), ultimate tensile strength (UTS), modulus of elasticity (E), and elongation percentage are illustrated in table 1. As expected, there is little difference between those plates with more than double laser power. Using both 0 degree plates as reference to be compared, 1070 W bars showed an average YS of 100 Ksi with a standard deviation of 1.08 Ksi and UTS of 159.4 Ksi with a 1.43 Ksi deviation while 2620 W bars showed a YS 97 Ksi and UTS 158.9 Ksi with standard deviation 0.39 Ksi and 2.09 Ksi320respectively.

In terms of ductility, the higher wattage provided a slightly higher ductility by elongating approximately 2.5% more than the original length even though ductility may be affected by lower density. However, as reported by Colton, et al., AM materials ductility is affected only if the material porosity is 5% or higher [1].SpecimenAverage Ultimate Tensile Strength (Ksi)Standard DeviationAverage 0.2% Yield Strength (Ksi)Standard DeviationAverage Modulus of Elasticity (Msi)Standard Deviation0 Degree 1070 W159.41.08100.731.438.411.5020 Degree 1070 W153.93.3896.91.719.833.2530 Degree 1070 W158.61.0298.382.428.250.830 Degree 2620 W158.90.3996.982.099.421.7920 Degree 2620 W156.33.494.82.6011.612.6630 Degree 2620 W155.31.1690.31.3112.41.77Table 1.

Results from tensile strength testingFurthermore, samples with the same laser power (1070 W) starting with 20-degree sample showed a 3% lower UTS and YS than sample with 0 degrees. However, it elongated 6% more than the corresponding sample. On the other hand, the 30-degree sample has a similar ultimate tensile strength, and yield strength to the 0-degree specimen, a difference close to 1% for both properties. Moreover, 30 degrees plates elongated 6% more than the 0-degree sample. Having similar tensile data can be attributed to the heat treatment homogenizing the microstructures for all power/angle combination.Low Cycle FatigueAs mentioned before, HR1 is used in applications such as cooling nozzles. Since they are part of the propulsion system of liquid engine rockets, they are exposed to high cyclic stresses. Therefore, the number of cycles required to failure was measured by reversal cycles to a 1% strain range and 0.5 Hz of frequency.

Therefore, the number of cycles required to failure was measured by reversal cycles to a 1% strain range and 0.5 Hz of frequency. As seen in table 2, the results for fatigue bars from 2620 W plates are graphically represented. Varying the angle from 0 to 20 degrees increased the number of cycles required to failure by 1100 reversal cycles. However, when the 30-degree sample withstood 5,500 cycles. In other words, there is a statistical difference between the number of cycles that 0- and 20-degree samples, but 30 degrees sample did not show to follow this trend. This difference may be a result of an error or variation during mechanical testing.321Average Elongation (%)37.843.844.335.332.341.0Standard Deviation (Ksi)2.530.851.552.637.630.96BSpecimen0 Degrees 10170 W 30 Degrees 1070 W 0 Degrees 2620 W 20 Degrees 2620 W 30 Degrees 2620 WNo of Reversal Cycles (Average of three samples) 4800 5100 4700 5800 5500Table 2.

Comparison of average number of reversal cycles needed to fail.FractographySEM was used to analyze the fracture surface along all different power/angle configuration. Figure 6 shows how one of the tensile bars’ surfaces that was selected randomly from 0-degree plate has fractured. A high number of gas entrapped pores were appreciated at the center of the surface (indicated with red arrows). In addition, dimples are also seen throughout the whole surface as proof of a ductile fracture mechanism. These dimples which indicate micro-void coalescence mechanism [4], were seen by increasing the magnification to 1000X. There was not a significant difference between all tensile specimens in terms of fracture surface. Fatigue bars were not included in this literature since they have not been analyzed under the scanning electron microscope yet.CFigure 6. A) Fractography images obtained using scanning electron microscope. B) Gas-entrapped pores found at surface.

A) Fractography images obtained using scanning electron microscope. B) Gas-entrapped pores found at surface. C) Dimples and cone-cup shape surface322AConclusionThis literature examined the defect content, microstructure, and mechanical properties of HR1 printed with Laser Powder Direct Energy Deposition at two different power settings, 1070 W and 2620 W and at three different deposition angles including 0, 20, and 30 degrees. Based on the results, the following conclusions have been derived:Increasing laser power increases the width and depth of melt pools, increasing the defective content on the material. However, the difference is not significant enough to affect ductility.

However, the difference is not significant enough to affect ductility. • NASA HR1 heat treatment was effective to homogenize the microstructure in all samples byproviding a similar austenitic microstructure with similar grain size and shape.Similar microstructure provided, as expected, similar mechanical properties in terms of tensile strength and low cycle fatigue having no statistical difference.Fractography proved that titanium segregation was not present in any of the samples due heat treatment homogenization since all specimens had dimples showing a ductile fracture mechanism present at fractured surfaces.REFERENCES[1] C. Katsarelis, P. Chen, P. Gradl, C. Protz, Z. Jones, Additive Manufacturing of NASA HR-1 Material for Liquid Rocket Engine Component Applications[2] A. Soltani-Tehrani, P. Chenc, C. Katsarelis, P. Gradl, S. Shaoa, N.

E EFFECTS OF DRY HEAT STERILIZATION ON PARTS USING SELECTIVE LASER SINTERING Mitchell J. George* and Richard H. Crawford† *Spectral MD, Inc., Dallas, TX 75390 Dept. of Mechanical Engineering, The University of Texas at Austin, Austin, TX 78712Abstract Selective Laser Sintering (SLS) is a manufacturing process that can build arbitrarily shaped parts without part specific tooling. Its advantages have been employed in many different fields, one of these being medical surgery. Currently, SLS is limited in medical applications as a pre-operative modeling tool. For SLS manufacturing to progress in areas like compliant surgical tooling and patient specific bone matrices, concurrent work is needed to investigate the effects of medical sterilization on SLS materials. This paper presents the results of sterilization experiments on SLS parts built from nylon 11.

This paper presents the results of sterilization experiments on SLS parts built from nylon 11. To simulate the process of introducing tools into a sterile environment, these specimens were subjected to multiple rounds of dry heat sterilization. Changes to the dimensions, tensile strength and flexibility were recorded and analyzed. It was found that the specimens’ dimensions remained relatively constant. Both the tensile modulus and the flexural modulus decreased as the sterilization cycles progressed. The tensile modulus decreased by 25% and the flexure modulus decreased by 19% after ten rounds of sterilization.IntroductionThe nexus between medicine and engineering grows stronger as the relational benefit thetwo shares is realized. Both fields exist for the betterment of humanity so it is natural that they have concurrent and contributing applications. medical research.

Both fields exist for the betterment of humanity so it is natural that they have concurrent and contributing applications. medical research. The goal of this research is to advance the applications of selective laser sintering (SLS) into manufacturing parts for sterile fields.Selective laser sintering is one of a number of manufacturing techniques broadlyclassified as Solid Freeform Fabrication (SFF). This area of manufacturing includes different layer-based manufacturing processes that have significant advantages over traditional manufacturing methods. SLS uses powdered materials to build parts without the spatial constraints that limit typical machining processes. A computer model (usually built using computer-aided design or CAD) is sliced into thin cross sections that are selectively melted (also known as sintering) according to the cross-sectional geometry using a computer controlled laser.

Because unsintered powder surrounds parts, there is no need for support structures for overhangs or complex geometries. A leveling roller spreads fresh powder over the powder bed, and the laser selectively melts, or sinters, an area of the powder based on the part’s cross-section. With advances in materials and processing technologies, SLS has found application in many fields of engineering and manufacturing. In particular, in the field of surgery and surgical reconstruction, SLS has emerged as an appropriate tool for more than just physical modeling [1].Since SLS selectively forms its parts by sintering powder using a laser, no tooling isrequired. Because of this, inventory is limited to providing sufficient powder quantity for builds. The machinist’s skills are unnecessary; a computer directs the SLS machine to produce parts with dimensions as precise as ±.005” [2].

The machinist’s skills are unnecessary; a computer directs the SLS machine to produce parts with dimensions as precise as ±.005” [2]. This direct data-to-part process has interesting implications, because parts could be potentially designed on one side of the world in CAD, emailed to a service bureau for manufacturing, and finally delivered to the end user. In fact, oneof the first investigative applications of SLS was to send a “three dimensional fax” in 1991 [3]. Also, since SLS parts are not limited spatially, a complex multiple part assembly can be manufactured with the parts in the assembled state.Potential ApplicationsSLS has found applications in many industries, one in particular being medicine. Thisresearch explored the use of SLS to create parts such as medical instruments, implantable devices or bone matrices.

Thisresearch explored the use of SLS to create parts such as medical instruments, implantable devices or bone matrices. Existing and potential applications of SLS within the field of medicine are discussed in this section.Implantable devices used in reconstructive surgery, such as hip or knee replacements, aretypically not patient specific. Building devices such as these orthopedic implants using SLS would allow unique geometries customized to the patient. Materials used in reconstructive surgery are limited by their biocompatibility. Hydroxyapatite is a biocompatible candidate appropriate for SLS that could be used for these kinds of applications [4]. Bone replacement therapy is practiced to combat traumatic injury or disease. One major problem with current therapies is rejection, since the human body is very particular with regard to bone microstructure.

One major problem with current therapies is rejection, since the human body is very particular with regard to bone microstructure. Since it is capable of producing small intricate geometries, SLS is a potential approach for constructing patient specific bone replacement. SLS has already been applied to the construction of patient specific prosthetic sockets [5]. Three-dimensional scans of a patient’s residual limb provide a foundation for the design of these sockets in CAD. Also, prosthetic feet and ankle-foot orthotics have been prototyped [6, 7].Research ObjectivesIt is hypothesized that dry heat sterilization will have no appreciable effect on thedimensional or material properties of nylon 11 parts manufactured with SLS. The SLS process requires temperatures up to 185° C, and the dry heat sterilization process used in this thesis only reaches a temperature of 121° C.

The SLS process requires temperatures up to 185° C, and the dry heat sterilization process used in this thesis only reaches a temperature of 121° C. Also, it is standard practice to use non-virgin powder (overflow or part cake powder) to build parts with minor loss of strength. For this reason, the nylon 11 material should not degrade due to dry heat sterilization. To prove this hypothesis, the following objectives were pursued:1. Build test specimens suited for flexural and tensile testing.2. Subject these specimens to multiple rounds of sterilization while removing random specimens in each round to create groups of appropriate sample size with varying sterilization exposures.3. Monitor the dimensional changes of each specimen following each sterilization cycle.4. Apply tensile and flexural tests to the specimen groups, paying close attention to the number of sterilization rounds associated with each group.5.

Apply tensile and flexural tests to the specimen groups, paying close attention to the number of sterilization rounds associated with each group.5. Observe and analyze any trends in the dimensional, flexural and tensile data.Experimental MethodsThe tests performed in this research were guided by pre-established standards. Effortswere made to reduce the variability of the testing procedures and increase consistency. Alltensile and flexure testing was performed on an Instron 3340 Series machine at The University of Texas at Austin. Similar tests were conducted on a United SSTM-20 kN tensile tester at Harvest Technologies (Belton, TX) to verify the data obtained UT Austin. Both platforms used external strain gauges to monitor elastic deformation in the tensile tests.The test methods used in this research were guided by American Society for Testing ofMaterials (ASTM) standards. The flexure testing standard is ASTM D790.

The flexure testing standard is ASTM D790. This method requires that the specimen being loaded to break or fail within the 5% strain limit of the test. Otherwise, accurate flexural strength cannot be determined. The tensile test standard is ASTM D638. This standard is designed to produce tensile property data for the specification of plastic materials.Each flexure specimen was placed on a three point bending platform, centered on the twobottom supports. Only one specimen was tested at a time. The flexure test began with the upper member touching the flexure bar, but causing no deflection. The tension test began with the tensile bar gripped in both the upper and lower jaws. Figure 1 shows one of the specimens in tension.Figure 1. Tensile test.To accurately record cross head displacement during non-plastic deformation, an extensometer was placed on the specimen until the program prompted the user to remove it.

Tensile test.To accurately record cross head displacement during non-plastic deformation, an extensometer was placed on the specimen until the program prompted the user to remove it. The extensometer is seen in Figure 1 attached to the middle of the tensile specimen.The test specimen types were chosen based on the material characteristic tests performedand the polyamide material used, nylon 11. The tensile specimens were specified by ISO 3167. Figure 2 shows a representative picture of the specimen and its dimensions. This type of specimen was chosen because it is commonly used for many different tests with plastics, specifically tensile testing.Tensile Specimen DimensionsOverall Length: 150 mm Thickness: 4 mmCenter Section Width: 10 mm Length: 80 mmFigure 2. ISO 3167 with dimensions.The flexure specimen specified by ASTM D790 is appropriate for the three point bendtest.

ISO 3167 with dimensions.The flexure specimen specified by ASTM D790 is appropriate for the three point bendtest. For thermoplastic or thermoset materials, D790 suggests that a rectangular specimen of dimensions 127x12.7x3.2 mm3 be used.The specimens used for testing in this research were built using a 3D Systems Vanguard HiQ Sinterstation. The build parameters for the test specimens are given in Table 1: The material used in the build was 100% virgin nylon 11 powder.Table 1. Description of Build Parameters Laser Power Scan Spacing Part Bed Temp. Feed Bin Temp.(Watts) 36(inches) 0.008(Celsius) 186(Celsius) 186A diagram describing the orientation of the test specimens during the build is shown in Figure 3. The layers in the figure provide a convenient means of describing how the build was set up, as discussed below.A BzLayer Key A: Flexure bars B: Tensile bars C: Other parts D: Flexure bars E: Tensile barsCxD EyBuild AreaFigure 3.

Diagram of Build Setup for Test Specimens.The build area outlined in Figure 3 describes the walls in the build bin in the SLS machine. The front face of the layers corresponds to the front face of the SLS machine, with the z direction defining the direction of height increase through the build process. Layer C contained parts built for other research projects. The test specimens for this research were located in layers A, B, D and E. Each of these layers contained two layers of the same type of specimen, either flexure or tensile.The entire build was removed from the machine 12 hours after the build was completed,and then cooled another 6 hours on a processing table. All parts were labeled as the break out progressed according to their height and location in the build. There were two parts discovered to have been built with deformities. Both parts were on the outer extreme of the build area.

There were two parts discovered to have been built with deformities. Both parts were on the outer extreme of the build area. It is possible that these parts cooled too fast and curled, since layers from these parts were found in the overflow bins. The powder roller most likely lifted these parts as they curled and pushed them into the overflow bins. An initial check of one flexure specimen gave a density of .9349 g/cm3. The published density of nylon 11 is 1.01 g/cm3.A Getinge (Rochester, NY) Castle 133 Vacuum Steam Sterilizer autoclave was used forthis research. The temperature and pressure specifics are presented in Table 2. The specimens were not stored in any special manner between cycles, and cycles were run every 4-5 days. To observe trends in material properties due to the effects of sterilization, groups of test specimens were cycled through from one to ten rounds of sterilization.

To observe trends in material properties due to the effects of sterilization, groups of test specimens were cycled through from one to ten rounds of sterilization. After all of the samples completed the first cycle, four tensile and four flexure samples were randomly chosen for removal. This process was repeated for ten rounds. After ten cycles, there were groups of specimens that had been sterilized once, twice and all the way up to ten times. A control group of ten flexure and ten tensile specimens was never put through the autoclave cycle.Table 2. Sterilization thermal and pressure cycle.Time (min:sec) Stage 0:00 Purge Condition 3:37 Heat Up 4:52 Exposure 6:23 26:23 Exhaust 27:07 Vacuum Air In 42:08 Complete 42:57Temp. (°F) 131 234 160 251 251 160 120 125Pressure (psig) 0.38 9.09 -19.83 16.77 16.57 -19.85 -26.12 -3.85After each sterilization cycle, the flexure samples were dimensioned using a set of digitalcalipers with an accuracy of ±.001”.

Each specimen’s length, width and height were recorded after each cycle until the specimen was removed from the sample population. The dimensions of the flexure bars were taken from the same points of contact each round. Width was taken from what the user perceived as the middle, and height was taken at the edge of the part at this same location. The tensile specimens were not dimensioned because of their awkward dog bone shape. The control group of tensile specimens was also dimensioned after every cycle and recorded with the test samples. This allowed the observation of any trends in volume change. It also provided a good indication of the validity of this data. The percent change in length, width and height were recorded. From this data, the percent change in volume of the specimens versus sterilization cycle was produced.ResultsFigure 4 shows the percent change in volume versus the sterilization cycle.

From this data, the percent change in volume of the specimens versus sterilization cycle was produced.ResultsFigure 4 shows the percent change in volume versus the sterilization cycle. There are 40data points for cycle 1 since there were 40 flex samples put through the first round of sterilization. The six controls do not contribute to this data since they were never sterilized. The average percent change was found for each cycle and graphed in a similar manner in Figure 5. A positive percent change means that, on average, the volume of the parts increases. A negative volume change, like in cycles 2, 5, 7, 9 and 10, indicates an average loss of volume in the specimens. The small variations in volume change are probably due to limited measuring resolution and not actual volume changes.% Change in Volume versus cycle43.2) 3 ^ m m1( e m u o Vl010123456789234cycleFigure 4.

Data corresponding to maximum load, maximum stress and flexure extension at maximum flexure load will not be discussed. The mean and standard deviation of flexural modulus for each cycle was obtained from the data. These values were calculated by the testing program.To investigate changes in test specimen flexibility, the flexural modulus of each groupwas graphed versus cycle number. This graph can be seen in Figure 7, along with error bars corresponding to the standard deviation of each cycle. It is clear that the flexural modulus decreases as the number of sterilization cycles progresses. This means that the average stress required to deflect the bars a constant amount decreases as the number of cycles progresses. In other words, the bars become more flexible and less resistant to bending as the number of cycles increases.

In other words, the bars become more flexible and less resistant to bending as the number of cycles increases. The average flexural modulus of the control specimens was 1,266 MPa, while the flexural modulus of after ten sterilization cycles was 1,026 Mpa, a 19% difference.It should be noted that the standard used for this test, ASTM D790, requires that the specimen either break or fail within the 5% strain limit of the test. None of the flexure specimens in this research broke when tested.Flex Modulus vs Rounds of Sterilization1400) a P M ( k l13501300) a p M ( s u u d o M x e l Fl12501200115011001050Modulus100095090001234567891011Round of SterilizationFigure 7. Flexural modulus versus sterilization cycle.The output data for the tensile specimens of cycle 1 is seen in Figure 8. There areonly four specimens in this cycle because the specimen clamps were not tightened during the tension test of specimen 3.

There areonly four specimens in this cycle because the specimen clamps were not tightened during the tension test of specimen 3. The output graphs display tensile stress versus specimen extension. The pre-yield slope of these curves produces the modulus. The yield points, indicated by small black triangles, are all relatively close while the extensions at break vary greatly. The break pattern seen in this specimen group is representative of the other ten groups.Similar to the data seen from the flexural specimens, the tensile modulus is presentedversus sterilization cycle in Figure 9. The sudden dip in elastic modulus during cycle 7 is most likely due to user error. Since four specimens represent each cycle, an anomaly of this magnitude is unlikely. It must be noted that the modulus recorded for cycle 7 is the lower than all following cycles.The data show that after 10 sterilization cycles, a modulus loss of 25% occurred.

It must be noted that the modulus recorded for cycle 7 is the lower than all following cycles.The data show that after 10 sterilization cycles, a modulus loss of 25% occurred. As themodulus decreases, less force is required to induce the same amount of specimen extension.Figure 8. Tensile test data for cycle #1.Tensile Modulus vs Rounds of Sterilization150014001300) a P M ( s u u d o Ml12001100mod100090080001234567891011Round of SterilizationFigure 9. Tensile Modulus versus sterilization cycle.The flexure modulus and tensile modulus of the specimens appear to degrade as sterilization cycles progress. To distinguish between actual material degradation and statistical variance, a Student’s t-test was performed. The test performed investigates the probability that the trends seen in the flexure and tensile moduli were due to material degradation and notstatistical variance.

The test performed investigates the probability that the trends seen in the flexure and tensile moduli were due to material degradation and notstatistical variance. Specifically, the t-test will indicate when the difference between the means of the control and test specimens is statistically significant. This test is influenced by the number of specimens in each sample and their standard deviations.For each case, the sample sizes for the control and test specimens were both 4. Based on the moduli and their standard deviations, a t-statistic was computed for each cycle. This number was compared to the corresponding t-distribution value, which is based on the desired level of certainty (95% in this case) and the degrees of freedom (6 in this case, 8 total data points in the two groups minus 2) [8]. If the calculated t-statistic is greater than the distribution value, then the two means are significantly different.

If the calculated t-statistic is greater than the distribution value, then the two means are significantly different. The critical value for 6 degrees of freedom and a 95% level of certainty is 2.447. This value is the same for both the tensile and flexural data.The results of the t-test for the flexure modulus are presented in Table 3. As the tableshows, the critical value is reached at cycle 5. At this point, the t-statistic exceeds the distribution value and the two means become statistically different.Table 3. T values for flexure modulus.Cycle T Statistic1 2 3 4 5 6 7 8 9 100.060009 1.24427801 1.3176711 2.32970326 3.15404731 2.81308329 4.48828551 3.15255105 3.97945406 4.57059951The results of the t-test for the tensile modulus are presented in Table 4. Only cycles 1and 3 exhibit means that are not significantly different from the control mean. The high t-statistic value from cycle 7 is due to the fact that its mean was significantly lower than the control mean.

The high t-statistic value from cycle 7 is due to the fact that its mean was significantly lower than the control mean. While the mean of cycle 8 was not as low as cycle 7, it had a very tight standard deviation and a high t-statistic was calculated. This t-test analysis confirms that the trends seen in the both flexure and tensile moduli indicate decreasing mechanical properties.Table 4. T values for tensile modulus. Cycle T Statistic1 2 3 4 5 6 7 8 9 101.247057299 2.749511866 2.108490685 2.833899243 2.628964314 3.148677132 6.711802395 6.306763569 3.18092205 2.962812139DiscussionThe changes in the flexure specimen dimensions were small relative to the precision ofthe dimensioning tool used. The caliper used to dimension all of the samples has a rated accuracy of 0.02 mm and a resolution of 0.01 mm. Because of this, differences of 0.02 mm between cycles are not significant and can be attributed to uncertainty. A standard flexure volume is 5556.25 mm3.

Because of this, differences of 0.02 mm between cycles are not significant and can be attributed to uncertainty. A standard flexure volume is 5556.25 mm3. The maximum error equivalent would be 5597.02 mm3. The difference between these volumes is 0.01%; therefore changes of volume of 0.01% are not significant. The actual changes of the specimen dimensions for width and height were on the order of 0.01 mm to 0.05 mm. For length, it was often on the order of 0.05 to 0.1 mm. Figures 4 and 5 include this error, and cycles 1-2 and 7-10 show volume changes above 0.01%. Also, there was no noticeable increase in standard deviation of the percent changes due to the decrease in the number of specimens as the cycles progressed.It should be noted that the appearance (color, shape and texture) of the flexure specimensnever changed throughout the sterilization cycle process. Interestingly, the ink from the permanent marker used to mark the parts bled through to nearby areas of the specimens.

Interestingly, the ink from the permanent marker used to mark the parts bled through to nearby areas of the specimens. Referring to Figures 4 and 5, the volume of the flexure bars seems to be non-constant. An increase in volume of 1.71% was recorded after cycle 1, and a decrease in overall volume of - 1.21% was recorded after cycle 10. It is difficult to prove the volumes of the specimens were actually ever changing, even after cycles 1 and 10, since there is no control comparison. It would have been beneficial, from an analysis standpoint, to measure a group of controls that were not being sterilized after the end of each cycle to compare. Also, the method of measuring the samples was assumed to be a constant process but in fact the points of contact made for measuring the specimens varied for each round.

Also, the method of measuring the samples was assumed to be a constant process but in fact the points of contact made for measuring the specimens varied for each round. With the variability in dimensioning and the inherent error in the measurement method, the volumes of the flexure specimens are determined to be unchanged by the sterilization cycle used in this research.Published values for the flexural modulus vary. Table 5 shows various modulus valuespublished by different companies and research groups.Table 5. Published values of the flexural modulus of Nylon 11.Flex Modulus Company/Group Name (psi) 126,000 Directed MFG 145,000 Indian Plastic Portal Polymer Web 145,000 Granta CES Edupack 133,000 162,000 Arkema 186,000 3D Systems(MPa) 869 999 999 917 1120 1285The value published for Arkema’s Nylon 11 is the relevant value in Table 5 since the material used in this research was purchased from Arkema.

The original modulus of the parts was 1266 MPa, while Arkema’s published value is 1120 MPa. The modulus of the flexural specimens did not drop to this value until cycle 4. The value of this modulus is less important to this research than its change over the course of the sterilization cycles. However, a possible explanation for this difference in modulus is that the flexure specimens did not break after the 5% strain required by ASTM D790. Also, differences in process parameters used to build the specimens could result in differences in the moduli.From Figure 7, it is evident that the average flexural modulus for each cycle decreases as the rounds of sterilization progress. Since the sterilization cycles exposed the flexure specimens to elevated heat without water vapor, the degradation of the parts can be attributed strictly to thermal effects. Unfortunately, there is generally no accepted mechanism for thermal decomposition of nylons.

Unfortunately, there is generally no accepted mechanism for thermal decomposition of nylons. There are many proposed chemical degradation mechanisms, but none of these occur until 300° C [9]. Since the exposure temperature during the sterilization cycles was 121° C, none of these established degradation mechanisms were can explain a loss in flexural modulus. Instead, degradation due to the complexities of the microstructure of SLS parts probably caused this drop in modulus. The microstructure of nylon parts built using SLS is atypical due to the layer-based manufacturing process and the inherent porosity. At this point, thermal degradation mechanisms in nylon parts built using SLS are unknown. One hypothesis for this mechanism is a decrease of molecular weight of the nylon particles, which causes shortening of the polymer chain lengths.

One hypothesis for this mechanism is a decrease of molecular weight of the nylon particles, which causes shortening of the polymer chain lengths. This in turn could reduce the crystallinity of the material [10].Similar to the flexural modulus, the tensile modulus of nylon 11 has manydifferent published values. Several published values can be seen in Table 6. Again, these values are provided for reference to the tensile specimens’ actual tensile strength, but the change in tensile strength is more important than the nominal value.Table 6. Published values for tensile modulus of nylon 11.Company/Group Name Directed MFG Granta CES Edupack 3D Systems Indian Plastic PortalTensile Modulus(kpsi) 202-275 180-190 232 261-268(Mpa) 1,392-1,896 1,240-1,310 1600 1,850-1,800The tensile modulus decreased more than the flexural modulus after 10 cycles ofsterilization (25% versus 19%). On the whole, the modulus decreased after every round.

On the whole, the modulus decreased after every round. The increases in rounds 3 and 8 are exceptions but the focus of this research is the general trend of decrease in modulus as the number of cycles increases. Similar to the discussion concerning flexural modulus degradation, reasons for the degradation of these tensile specimens are difficult to find.ConclusionsThe losses in flexural and tensile modulus due to 10 rounds of dry heat sterilization aresignificant. Also, it is assumed that the trends observed in Figures 7 and 9 will continue as sterilization cycles continue. This research was a first attempt at quantifying this phenomenon, and with repeated experiments an average percent loss of material integrity per cycle can be established. This implies that for certain applications, if a required factor of safety is known and the material properties of a certain part are known, a maximum number of times a part can be safely sterilized can be determined.

For example, a surgical tool built by SLS will be designed with a certain factor of safety to ensure the tool will not fail due to the strength demands of the surgical procedure. As the tool is sterilized and reused for different procedures, this factor of safety will decrease. Based the results of this research, the rate at which this safety factor decreases can be estimated and tool life for specific applications can be established. It should be noted that dry heat sterilization is not the accepted method for sterilization of critical items, which require steam sterilization. Additional effects from normal steam sterilization could further degrade sintered parts. This is a topic for future research.References 1 Janssens, M., and Poukens, J., “Rapid Technologies in Medicine: What Can, and Can’tBe Done and Why”, Materialise N. V., January 31, 2007.2 Solid Concepts Incorporated.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEffects of Electric Field on Selective Laser Sintering of Yttria-Stabilized Zirconia Ceramic PowderDeborah Hagen, Desiderio Kovar and Joseph J. BeamanDepartment of Mechanical Engineering University of Texas at Austin Austin, TXSolid Freeform Fabrication Symposium Austin, Texas, USA August 13-15, 2018AbstractSelective laser sintering (SLS) of ceramic material is particularly challenging. High sintering temperatures and slow sintering kinetics of ceramic material combined with poor thermal shock resistance have resulted in cracking when ceramics are sintered to full density one layer at a time. This work investigates the use of an electric field applied simultaneously with laser scanning to accelerate the kinetics of sintering to produce a multi-layer SLS ceramic part.

This work investigates the use of an electric field applied simultaneously with laser scanning to accelerate the kinetics of sintering to produce a multi-layer SLS ceramic part. Ceramic sintering rates have been shown to increase by orders of magnitude during conventional furnace-based flash sintering, in which electric field applied simultaneously with furnace heating. In this work, we investigate the effects of an electric field applied during SLS processing of yttria-stabilized zirconia ceramic.Additive ManufacturingAdditive manufacturing (AM) is a layer-based manufacturing method that is capable of making a part directly from a computer-aided-design (CAD) file. Additive manufacturing does not require expensive part-specific tooling, which drastically shortens the time between design and the build of a part as well as reducing the cost due to tooling.

Additive manufacturing does not require expensive part-specific tooling, which drastically shortens the time between design and the build of a part as well as reducing the cost due to tooling. Most AM methods also can be utilized to build intricate and complex-shaped parts which are difficult or impossible to build with traditional manufacturing methods.Selective laser sintering (SLS) is a type of AM in which a high-powered laser fusespowdered materials. Typically, a layer of powder is applied to the top of the build chamber. A laser is then focused onto a layer of powder that sits on top of the build chamber with a set of optics. A pair of mirrors rotate so that the reflected laser beam rasters across the top layer of powder in a pattern that melts or sinters only the powder that will become part of the solid formed shape.

A pair of mirrors rotate so that the reflected laser beam rasters across the top layer of powder in a pattern that melts or sinters only the powder that will become part of the solid formed shape. After the layer is lased, the piston drops the build chamber slightly and another layer of powder is applied to the top surface and this layer is then scanned with the laser. The layer application and selective laser scanning steps are repeated until the entire part is completed.[1]Complex shapes and detailed parts are achievable using SLS processes. Supports are notnecessary for overhanging or sloped surfaces as they are for stereolithography (SLA)[2] and909fused deposition modelling AM methods because the powder bed is self-supporting.

Supports are notnecessary for overhanging or sloped surfaces as they are for stereolithography (SLA)[2] and909fused deposition modelling AM methods because the powder bed is self-supporting. Low temperature polymers, high temperature polymers, and metals are being commercially produced using a direct SLS process.Additive Manufacturing of CeramicsCeramic additive manufacturing has been primarily focused on indirect sintering methods that utilize a binder or adhesive to fuse ceramic powder particles together and create a solid part. In contrast, direct sintering does not form the shape of a part using adhesives or binders. Instead, direct sintering uses energy source(s) to fuse powder particles directly together when forming the shape of the part. SLA[2], binder jet printing[3], and extrusion-based ceramic AM processes[4] all use polymers to temporarily fuse ceramic powders together into the desired shapes.

SLA[2], binder jet printing[3], and extrusion-based ceramic AM processes[4] all use polymers to temporarily fuse ceramic powders together into the desired shapes. The polymers that are used for shape formation for all of these ceramic AM processes is subsequently removed by pyrolysis, which can fracture the part unless it is performed very slowly. The time required for crack-free pyrolysis increases exponentially with part thickness, essentially limiting the size of ceramic parts that can be economically produced.[5] After the pyrolysis step, the ceramic parts are densified by a furnace firing.Attempts at direct SLS sintering processes of ceramics with little or no binder have hadlimited success.[6-8] These processes aim to provide enough energy to the process to completely sinter together ceramic powder particles one layer at a time. The intent is to fully densify each layer either by melting or sintering the ceramic.

The intent is to fully densify each layer either by melting or sintering the ceramic. The direct sintering approach is particularly challenging because the laser energy densities required to melt or sinter ceramics in the time it takes the laser to scan can be also be high enough to thermally shock the densified ceramic. In addition, shrinkage of approximately 50% by volume is required to densify a layer and such large shrinkages cannot be easily accommodated by the underlying layers of the brittle ceramic material.Laser Flash Sintering HypothesisWe propose an alternative approach to AM of ceramic parts in which a laser is used inconjunction with an electric field to induce initial stage sintering of the part to a sufficient degree that necks are produced between the particles such that the part will hold together wherever the laser was scanned over the powder bed. This process is repeated one layer at a time until a bulk part is fabricated using this layer-by-layer approach.

This process is repeated one layer at a time until a bulk part is fabricated using this layer-by-layer approach. Final densification is accomplished in a post-processing firing.Flash Sintering: Effects of Electric Fields on Sintering of Ceramics in Furnace EnvironmentsElectric-field assisted sintering in conventional furnace-fired ceramic processing hasexperienced a rapid increase in interest as the result of its ability to dramatically lower sintering temperatures and increase sintering rates compared to conventional sintering. Two types of electric field-assisted sintering experiments have been demonstrated 1) spark plasma sintering where sintering occurs within a die at high pressure[9], and 2) flash sintering where the field is applied to an otherwise bare specimen using electrodes.

Although the precise mechanisms that910cause the electric field to increase sintering kinetics during flash sintering remains controversial, it has been demonstrated that sintering temperatures can be hundreds of degrees lower than for conventional sintering and sintering times can be reduced from hours to seconds.[10-15]Laser Flash Sintering ProcessLaser flash sintering is an approach to AM of ceramic parts in which a laser is first usedin conjunction with an electric field to induce initial stage sintering of the part to a sufficient degree that necks are produced between the particles such that the part will hold together wherever the laser was scanned over the powder bed. The sample is then over coated with fresh powder and the laser is scanned over the surface again. This process is repeated until a bulk part is fabricated using this layer-by-layer approach.

The sample is then over coated with fresh powder and the laser is scanned over the surface again. This process is repeated until a bulk part is fabricated using this layer-by-layer approach. In the second step, the partially sintered part is removed from the powder bed, the loose powder is gently removed to expose the partially sintered part, and a post-process, conventional sintering treatment in a furnace or a conventional flash sintering treatment is performed to complete densification.Figure 1: Laser flash sintering process flowA DC electric field is applied to the powder bed using two electrodes placed on opposite sides of the build chamber. Powder bed layers can be applied as a slurry consisting of powder, a solvent, and a binder. After evaporation of the solvent, the remaining binder is used to temporarily hold the nano-scale ceramic powder in a dense layer, and a dispersant keeps the powder from forming pore-inducing agglomerates.

The binder breaks down thermally several hundred degrees lower than the ceramic sintering temperature and pyrolyses during the lasing process. A carbon dioxide laser beam is focused through zinc selenide optics and rastered across the powder bed with a pair of mirrors whose motion is controlled with precision galvanometers.Laser Flash Sintering ResultsThe literature on flash sintering reports experimental results performed in a furnaceenvironment; there are no previous reports of inducing flash sintering using a scanning laser heating rather than uniform furnace heating. Experiments with 8% yttria stabilized zirconia ceramic reported here demonstrate that flash sintering effects also occur during laser heating.Pressed green ceramic pellets of 8% yttria stabilized zirconia (Tosoh TZ-8YS), one inch in diameter and approximately 0.12 inches thick, were lased with and without the application of an electric field.

A DC voltage of 315 V was applied across electrodes painted onto the surface of the pellet with colloidal graphite paint. The gaps between electrodes was varied to give field strengths up to 900 V/cm by varying the distance between electrodes. The laser beam was911rastered in a serpentine patterned on the powder bed, with the edges of the serpentine touching both electrodes. Beam speed was 1 m/s, with at least three repetitions at various beam powers. The scan was done only on the surface of the pressed green pellet.The samples lased with electric field showed increased grain growth and more rapiddensification as compared to the samples lased in the absence of an electric field. Very bright light was emitted during lasing with an applied electric field, while a dull red glow was the only visible effect on the pellets lased without an electric field. An electric current was also measured during lasing when an electric field was applied.

An electric current was also measured during lasing when an electric field was applied. Electrical current is reported in the literature for flash sintering where electrodes are connected directly to the sample. The bright light emissions, the rapid densification (and subsequent over-sintering), and the electrical current demonstrate a flash effect when ceramics that are directly connected to a DC power supply are simultaneously lased. Figure 2 shows differences between the field and no-field conditions.During LFSMicrostructure of SurfaceWith Electric FieldWithout Electric FieldFigure 2: Comparison of SLS of yttria-stabilized zirconia pellets with and without electric field.

Visual appearance during scanning is shown on the left, and microstructure of the surface of the pellets in shown on the right.Future Work and ConclusionsBright light was observed during lasing of ceramic powder under an electric field thatwere not visible in the absence of an electric field, confirming that a scanning laser can be used to initiate flash sintering. Future work will investigate the effects of electric field on the processing of ceramic multi-layer parts with SLS process. Application of multiple ceramic powder layers and adhesion of these layers to one another through direct sintering will be investigated. The effects of lasing and electric field parameters on shape formation, microstructure, and defects will be examined.

viewed PaperEffects of Fast Neutron Irradiation on the Microhardness of Inconel 625 and Inconel 718 Fabricated via Laser Powder Bed FusionM. Andurkar1, V. O'Donnell2, T. Keya3, B. C. Prorok3, J. Gahl2, S. M. Thompson*11 Alan Levin Department of Mechanical and Nuclear Engineering, Kansas State University, Manhattan, KS 66502, United States of America.2 University of Missouri Research Reactor (MURR), University of Missouri, Columbia, MO 65211, United States of America.3 Department of Materials Engineering, Auburn University, Auburn, AL 36849, United States of America.Corresponding author: smthompson@ksu.eduAbstractThe demand for advanced materials in constructing next generation nuclear reactors has intensified the need to explore additive manufacturing (AM) processes as an alternate means of fabricating components. In this study, Inconel 625 (IN625) and Inconel 718 (IN718) samples fabricated using Laser Powder Bed Fusion (L-PBF) were irradiated using fast neutrons.

In this study, Inconel 625 (IN625) and Inconel 718 (IN718) samples fabricated using Laser Powder Bed Fusion (L-PBF) were irradiated using fast neutrons. Samples investigated included as-printed and heat-treated at either 700, 900, or 1050 for 1 hour to understand the impact of heat treatment on any neutron irradiation hardening. Wrought IN625 and IN718 samples were also inspected for experimental control. All samples ℃ neutrons/cm2. To quantify were irradiated for 7 weeks resulting in a total fluence 2.74x radiation damage, the Vickers microhardness was measured before and after fast neutron irradiation. Results show that the IN718 samples experienced less change (-2.5 to 3.24 %) in microhardness. On the other hand, IN625 samples underwent more (0.9 to 7.21%) change in microhardness post fast neutron irradiation.1510Keywords: Selective Laser Melting, Nickel Superalloy, Vickers Microhardness, Radiation Embrittlement.1.

IntroductionThe Generation IV International Forum (GIF) which represent ten nations was founded in 2000 to develop six reactor technologies that can shape future, next-generation nuclear reactors [1]. The technologies are based on meeting the increasing energy demands on a sustainable level by building clean, safe, and cost-effective nuclear reactors. Most of these six technologies use closed fuel cycle aiming to make maximum use of fuel resources and minimize radioactive wastes. Out of these six technologies, three are planned to be fast reactors which employs fast neutrons for fission reactions. Neutrons produced during fission reactions with energy more than 0.1 MeV are known as fast neutrons. Fast reactors employing fast neutrons can provide significant advantage over the current slow or thermal neutron (E < 0.025 eV) reactors [2]. In fast neutron reactors, natural uranium is burned 60 times more efficiently than a normal reactor.

In fast neutron reactors, natural uranium is burned 60 times more efficiently than a normal reactor. Natural uranium (0.7% U-235 and 99.3% U-238) is burned directly by fast neutrons without it being converted to plutonium isotopes eventually utilizing uranium at maximum level. This will help to reduce the scarcity of natural uranium all over the world. Moreover, fast reactors have smaller core and higher power density compared to normal603reactors. Even though, fast reactors are researched to provide significant advantages over current nuclear reactors, there is a shortage of commercial and research fast reactors in the USA.To build safe, reliable structural components for high energy fast reactors, different material classes are qualified such as refractory metal alloys, Ni-based superalloys, Ti-based alloys. However, fabricating complex-shaped structural components using conventional machining methods is generally challenging due to high toughness and residual stress.

However, fabricating complex-shaped structural components using conventional machining methods is generally challenging due to high toughness and residual stress. To mitigate this issue, additive manufacturing (AM) technology serves as a novel, easier and more efficient way of building components. AM provides advantages over conventional machining like reduction in waste materials as AM operates on adding layers of material together to complete the final part in contrast to subtractive nature in conventional machining. Laser powder bed fusion (L-PBF) is one of the seven AM process categories specified in ASTM52900-21 [3]. LPBF process involves selective melting of material layer of 20-60 micron-thick within an enclosed setting of inert gas. Laser and powder interaction creates various byproducts like plume, spatter, powder ejecta due to surface tension and Marangoni effects [4].

Laser and powder interaction creates various byproducts like plume, spatter, powder ejecta due to surface tension and Marangoni effects [4]. Inert gas is used to efficiently carry these byproducts away from melt pool to obtain a clean and desired microstructure. It also reduces the oxidation of alloy and help preserve the chemical composition of alloy [5].Ni-superalloys are proposed as attractive material for building components to be used in GEN IV reactors [6]. IN718 and IN625 finds extensive application in structural components in thermal and fast reactors. Although Ni-superalloys have been successfully adopted in nuclear reactors for many years, research has confirmed that they are prone to become highly radioactive due to high neutron absorption cross-section. Instead of this drawback, IN718 finds its use in building tensioning springs, fuel channel spacers.

Instead of this drawback, IN718 finds its use in building tensioning springs, fuel channel spacers. IN718 is an attractive material in nuclear industry due to its properties such as good corrosion resistance, high strength due to the formation of phases like metastable phase at higher temperatures [7].’’ phase at lower temperature and equilibrium𝛾𝛾Nickel based superalloys are recommended to be employed in GEN IV reactors [6]. IN625 possesses excellent high temperature strength, creep, and toughness. It is been recommended in fabricating control rod and reactor core applications [8]. IN625 is a solid solution strengthening alloy due to the pronounced micro-segregation of elements like Mo and Nb during solidification. IN625 is becoming more ubiquitous in various industries like aerospace, nuclear or any other industries where harsh environments are expected [9].

IN625 is becoming more ubiquitous in various industries like aerospace, nuclear or any other industries where harsh environments are expected [9]. Conventional machining methods often face difficulties in processing IN625 components due to its high hardness, reduced machinability, and low thermal conductivity [10]. Fang et al. studied the microstructural evolution of IN625 specimens built via L-PBF subjected to different heat treatments [11]. It was found that as-built specimens possessed higher hardness due to high dislocation density arising from high thermal gradients during the L-PBF process. The microhardness values decreased when the as-built specimens were subjected to heat treatment temperatures at 870, 980, and 1150 oC for 1 hour. The samples were furnace cooled to room temperature. The heat treatment reduced the dislocation density and increased homogenization of grains.

The samples were furnace cooled to room temperature. The heat treatment reduced the dislocation density and increased homogenization of grains. On the other hand, IN718 is a precipitation hardening nickel-based superalloy which derives its strengthening from D022-ordered (Ni3Nb). Tucho et al. found that solution heat- treating IN718 samples (1100 oC for 1 hr) weakened the sub-grain boundaries indicating a 𝛾𝛾” reduction in dislocation density and micro-segregation of refractory metals like Nb and Mo. This ultimately resulted in a decrease of hardness compared to as-printed samples [12]. By performing heat treatment at different temperatures, different mechanical properties can be obtained from IN625 and IN718 alloys suitable for certain applications.604𝛿𝛿In the past, researchers have studied the effects of neutron and ion irradiation in IN718.

Byun and Farrell studied the changes in tensile properties of wrought IN718 in solution- annealed (SA) and precipitation hardened (PH) conditions. They found the SA IN718 showed a threefold increase in yield strength and a decrease in ductility. On the other hand, the originally stronger PH IN718 showed minimal radiation hardening post irradiation [13]. Because of high cost, long time, and radioactivity of neutron irradiated samples, researchers turned towards using ion irradiation to simulate the neutron irradiation damage. Hunn et al. studied the three different ions (Fe, H, He) irradiation effects in wrought SA IN718. They found out all three ions irradiation induced radiation hardening due to the formation of helium bubbles [14]. However, the irradiation damage depth from ion irradiation is very limited to the thickness in scales of nanometer to micrometers. This is not sufficient for measuring damage via microhardness testing [15]. Andurkar et al.

This is not sufficient for measuring damage via microhardness testing [15]. Andurkar et al. performed thermal neutron irradiation in L-PBF and wrought IN625 samples. They observed L-PBF built samples were lesser prone to detrimental radiation induced hardening compared to wrought samples [16]. Keya et al. conducted similar studies but using fast neutrons radiation in L-PBF and wrought IN625. From the preliminary results, they observed the L-PBF samples experienced both radiation induced hardening and softening for different fluence values [17]. Claudson studied the effects of neutron irradiation, with exposures exceeding (E >1 MeV), on the elevated temperature mechanical properties of wrought nickel-based and refractory metal alloys.

Claudson studied the effects of neutron irradiation, with exposures exceeding (E >1 MeV), on the elevated temperature mechanical properties of wrought nickel-based and refractory metal alloys. The results showed a decrease in ductility at elevated temperatures compared to those measured at room temperature [18].Although there have been several studies conducted on studying the irradiation effects in wrought IN718 and IN625, similar investigations in AM IN718 are scarce. There is a need of understanding on how AM built IN718 and IN625 would behave under similar radiation environment. For successful adoption of AM metals in nuclear industry, research studies can prove as an important guide on how to tailor the process parameters to obtain radiation hardening resistant component. This paper focuses on performing fast neutron irradiation on as-built, heat-treated, and wrought IN718 and IN625 samples.

This paper focuses on performing fast neutron irradiation on as-built, heat-treated, and wrought IN718 and IN625 samples. Vickers microhardness pre and post 5 weeks of fast neutron irradiation was performed to comprehend which sample condition was less prone to post irradiation hardening defect.2. MethodsGas atomized, virgin IN718 and IN625 powder with a mean particle size of 35 µm and 31.31 µm, respectively, were supplied by Carpenter Technologies. The elemental composition of Inconel 718 powder used in L-PBF in this study, following the ASTM F3055-14 standard for Additive Manufacturing Nickel Alloy (UNS N07718) with Powder Bed Fusion, is presented in Table 1. The chemical composition of Inconel 625 powder used for L-PBF process in this study adhering to ASTM Standard for Additively Manufacturing Nickel Alloy F3056-14 (21) is presented in Table 2. L-PBF Inconel 718 samples of dimensions 10 x 4 x 3 mm3 were fabricated using a Concept Laser Mlab 100R system.

L-PBF Inconel 718 samples of dimensions 10 x 4 x 3 mm3 were fabricated using a Concept Laser Mlab 100R system. The Concept Laser is equipped with a 100 W fiber laser with wavelength of 70 nm. The building process was carried out in an enclosed argon atmosphere for minimization of oxidization. Samples as shown in Fig. 1(a) were iteratively built at various process parameters and then inspected for density. Final process parameters consisting of power = 90 W, scan speed = 800 mm/s, laser diameter = 80 µm, hatch spacing = 60 m yielded 99.88% dense samples. For baseline purpose, wrought IN718 plate was procured from Metalmen NY, USA. The plate was manufactured as per ASTM B 670-07 (2013). Small 7.5 x 5 x 3 mm3 samples were sectioned from the plate using wire electrical discharge machining (EDM). The chemical composition of all forms of Inconel 718 samples used in this study is indicated in Table 1.

The chemical composition of all forms of Inconel 718 samples used in this study is indicated in Table 1. Inconel 625 cubes of 10 mm3 dimension were additively manufactured using same Concept Laser Mlab 100Rm, and layer thickness = 30𝜇𝜇𝜇𝜇605system. The samples were printed in an enclosed argon atmosphere for minimization of oxidization. Cubes were iteratively built at various process parameters and then inspected for density. Final process parameters provided for 99.79% dense samples, these process parameters were: power = 90 W, scan speed = 800 mm/s, laser diameter = 80 µm, hatch spacing = 60 µm, and a layer thickness = 25 µm. In this study, only vertically printed IN625 sample highlighted in Fig. 1(b) was used. Inconel 625 plate (152.4 x 152.4 x 4.76 mm3) was procured from Metalmen, NY, USA. The plate was manufactured per specification AMS5599. Small 1 x 1 x 0.5 cm samples were sectioned from the plate using wire electrical discharge machining (EDM).

Highlighted samples indicatethe optimized dense sample used in this study.The scan strategy used in this study is island scanning strategy with a checkerboard pattern with 5 × 5 mm squares. Each island alternated the scan direction perpendicular with- respect-to adjacent islands. Samples were separated from the build plate via wire EDM and then subjected to various heat treatments. The samples were subjected to three heat-treatment606temperatures of 700, 900, and 1050 ℃ for 1 hour. The rationale behind selecting these temperatures was the formation of phase in additively manufactured IN718 and IN625 as discussed in TTT diagram in a study conducted at by R. G. Thompson and National Institute of Standards and Technology (NIST) respectively [19][20]. Wrought samples were subjected to the same heat-treatment temperatures for 40 hours. Some L-PBF specimens were not heat treated, i.e., left in the as-built condition. Heating ramps used during heat treatment were ~5 ºC/min.

Some L-PBF specimens were not heat treated, i.e., left in the as-built condition. Heating ramps used during heat treatment were ~5 ºC/min. All specimens were subjected to air cooling to room temperature while remaining in the furnace. Uncertainty in heat treatment temperatures is ± 5 ºC.𝛿𝛿Fast neutron irradiation was carried out in cyclotron where a proton is accelerated at very high velocity using the electrically charged ‘dees’. The dees are connected to oscillating current source. The dees are sandwiched between two magnets which provides magnetic force. The proton with alternative electric and magnetic force gains high acceleration. It moves in concentric circles and exits once desired acceleration is achieved. The accelerated proton was bombarded on fluorine-18 target. Once the proton is bombarded it gives off energetic neutrons as a by-product following the 18O (p, n) 18F reaction [21]. The process is explained schematically in Figure 2.

Once the proton is bombarded it gives off energetic neutrons as a by-product following the 18O (p, n) 18F reaction [21]. The process is explained schematically in Figure 2. All IN625 and IN718 samples were irradiated for a duration of 7 weeks resulting in a total neutron fluence of 2.74 xneutrons/.10152 cmFigure 2: Schematic of fast neutrons production in cyclotron.Hardness is an indication of yield strength. Vickers microhardness test was performed using a Phase II 900-391D microhardness tester. On each sample, 5 measurements were taken before and after irradiation to be able to calculate average and standard deviation in measurements. Error bars on the plots represent the standard deviation. The indentations were made using 1 kgf (9.8 N) load applied with a 15 secs dwell time.3. ResultsThe Vickers microhardness of the pre-irradiated L-PBF as-built, heat treated and wrought IN718 and IN625 samples are shown in Fig. 3.

ResultsThe Vickers microhardness of the pre-irradiated L-PBF as-built, heat treated and wrought IN718 and IN625 samples are shown in Fig. 3. The as-built IN718 and IN625 samples measured 326.64 9.22 HV, respectively. This indicates the process parameters used for fabricating both materials did not have a major influence on the hardness of the samples. However, the slight variation in microhardness may potentially be due to difference in chemical composition of two materials. As-built IN718 and IN625 samples showed higher microhardness compared to their as-received wrought counterparts which had a microhardness of 232.42 ± 4.19 and 244.44 ± 5.73 HV respectively. The higher microhardness in L-PBF samples may be explained by them having fine dendritic6.21 and 327.26±±607microstructure and high dislocation densities arising from high thermal gradients present during the L-PBF process.

These high thermal gradients result from very high cooling rates (~105 K/s) and high heat flux imposed by the laser. The high cooling and solidification rates contribute immensely to trap the heavy atoms like Molybdenum (Mo) and Niobium (Nb) along interdendritic boundaries. These heavy atoms located along sub domain boundaries restrict the sliding of dislocations under external stress which ultimately increases the hardness of the material. The lower hardness in the as-received wrought samples is due to them possessing more homogenous and equiaxed grains microstructure due to prolonged heat treatment, lesser dislocation density. Lesser segregation of heavy atoms like Mo and Nb along grain boundaries since amount of Nb in wrought samples used in this study is lesser than the L-PBF powder. Moreover, the standard deviation in hardness of wrought samples is lower compared to L-PBF samples indicating more homogenous microstructure which ultimately yields homogenous property.

Moreover, the standard deviation in hardness of wrought samples is lower compared to L-PBF samples indicating more homogenous microstructure which ultimately yields homogenous property. Post three temperatures heat treatments, the microhardness trend in both materials showed similar trend.After the 700-1h heat treatment, the microhardness increased significantly for both materials. IN718 and IN625 measured 435.96 ± 12 HV and 353.46 ± 5.77 HV, respectively. The spike in microhardness may potentially be from the precipitation of intermetallic phase (Ni3Nb). This precipitate impedes the motion of plastic strain caused by the Vickers indenter 𝛾𝛾” during the hardness testing. The increase in microhardness in IN718 post 700-1h heat treatment was significantly higher than in IN625. This is due to higher quantity of Nb and Ta in IN718 compared to that in IN625 which may result in higher precipitation of phase.

This is due to higher quantity of Nb and Ta in IN718 compared to that in IN625 which may result in higher precipitation of phase. Increasing the heat treatment temperature from 700 , the microhardness of both materials decreased continuously. At 900 the microhardness of IN718 and IN625 measured ℃ , 347.18 ± 12 HV and 300 ± 6.6 HV respectively. Further increasing the temperature to 1050 the microhardness of IN718 and IN625 measured 277.2 ± 3.57 HV and 251.8 ± 6 HV respectively. Interestingly, L-PBF samples heat treated at 1050 ℃ reduced their hardness values and brought in the range of their as-received wrought counterparts microhardness values. The reduction in microhardness after heat treatment is possibly from elimination of dislocations, solid state transformations, grain recrystallization or grain coarsening, reduction in texture, and micro stress relaxation.

Similar observations have been reported [11], [22].to 900and 1050𝛾𝛾”℃℃℃,Figure 3: Microhardness of pre-irradiated as built, three heat-treated, and wrought IN625 and IN718 samples.608℃The microhardness values of L-PBF and wrought IN718 and IN625 samples after fast neutron irradiation are presented in Figures 4 and 5, respectively. It can be observed that fast neutron irradiation introduced radiation hardening/embrittlement in almost all samples. Helium is primarily produced in Ni-base alloys by high energy (fast) neutron reactions in natural Ni containing 68% 58Ni and 26% 60Ni, with the fast neutron n, α reactions. This can be explained by the neutron reaction presented in Equations (1) and (2) [23].

This can be explained by the neutron reaction presented in Equations (1) and (2) [23]. Helium bubbles are considered to embrittle grain boundaries by the stress developed growth of voids along the grain boundaries [8].58Ni + nf → 55Fe + 4He (1)60Ni + nf → 57Fe + 4He (2)Figure 4: Pre and post fast neutron irradiation microhardness measurements of all IN718 L- PBF and wrought samples. Error bars represent standard deviation in measurements.Figure 5: Pre and post fast neutron irradiation microhardness measurements of all IN625 L- PBF and wrought samples. Error bars represent standard deviation in measurements.609Although the change in hardness values in all samples is minimal post irradiation, it can result in serious damages like crack formation and ultimately result in fracture of critical end-user components. Irradiation-induced hardening was somewhat similar in all IN718 samples including wrought.

Irradiation-induced hardening was somewhat similar in all IN718 samples including wrought. This indicates that the structural integrity of AM parts under intense neutron irradiation is similar to that of conventional wrought samples. In contrast, fast neutron irradiation hardening had more influence on all IN625 samples. The percent change in microhardness post irradiation in all IN718 and IN625 samples are presented in Figure 6(a) and 6(b), respectively. IN718 was more resistant to radiation hardening possibly from the reason that IN718 have more Nb and Ta strengthening elements which could mitigate the fast neutron irradiation impact. Also, IN718 is generally known for its high strength at high temperatures compared to IN625 which is known more for corrosion resistance [24]. Similar radiation hardening or embrittlement in metals have been reported in past.

Similar radiation hardening or embrittlement in metals have been reported in past. [13], [25], [26].a) b)Figure 6: Percent change in microhardness of all a) IN718 and b) IN625 samples post fast neutron irradiation.4. ConclusionsNearly-full density IN718 and IN625 samples were fabricated using the L-PBF additive manufacturing process. The effects of fast neutron irradiation on as-built, heat-treated, and wrought conditions were studied. Vickers microhardness testing method was used to measure radiation-induced hardening in all samples.

Vickers microhardness testing method was used to measure radiation-induced hardening in all samples. Conclusions are summarized below:The as-built IN718 and IN625 samples had higher pre irradiation hardness value compared to their as-received wrought counterparts.Post heat-treatment, both materials showed similar trend in microhardness change with .Post fast neutron irradiation, all IN718 samples showed minimal percent change in ℃However, IN625 samples underwent more hardening with percent change inmicrohardness ranging between 0.9 to 7.21 %. IN718 samples displayed more resistance to radiation hardening (lower percent change) as compared to IN625 samples (higher percent change).The results provided in this study signifies the importance of adopting accelerated testing method to emulate the irradiation effects in materials considered important in building future efficient fast reactors.

The comparison of AM results along with currently used wrought610components will help boost confidence among designers/engineers to consider AM built components in future fast reactors. Fast neutron irradiation testing provides an accelerated mean of obtaining radiation effects without making materials radioactive which are difficult to handle. Different conditioned samples investigated in this study under same irradiation environment provides an insight into what condition samples performed superior to the others. The research results discussed in this study can provide assurance to certify and apply additively manufactured components in nuclear industry.AcknowledgementThis material is based upon work supported by the U.S. Department of Energy’s Officeof Nuclear Energy under Award Number DE-NE0008865.DisclaimerThis report was prepared as an account of work sponsored by an agency of the United States Government.

Department of Energy’s Officeof Nuclear Energy under Award Number DE-NE0008865.DisclaimerThis report was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights. Reference herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof. The views and opinions of authors expressed herein do not necessarily state or reflect those of the United States Government or any agency thereof.References[1] K. L.

E EFFECTS OF HEAT BALANCE ON THE VOID FORMATION WITHIN MARAGE 300 PROCESSED BY SELECTIVE LASER MELTINGT. Burkert*, A. Fischer**BMW Group, Rapid Technologies Center, Knorrstraße 147, 80788 Munich, Germany ** University Duisburg-Essen, Material Science and Engineering, Lotharstraße 1, 47057 Duisburg, GermanyAbstractThis contribution shows the results of a study that was conducted on the effects of varying selective laser melting (SLM) process parameters on the formation of microstructural voids within a maraging steel (type: Marage 300, 1.2709, AMS6514). Due to the large number of process variables the most influential parameters were identified first. These were energy input resulting from scanning speed, hatch distance, and layer thickness as well as the preheating of the platform. On the basis of the variation of these parameters the most abundant voids were identified and characterized by metallographic investigations.

On the basis of the variation of these parameters the most abundant voids were identified and characterized by metallographic investigations. Subsequently tensile tests derived information about the influence of such voids on the mechanical properties. Based on these analyses the reasons for the generation of such voids are discussed followed by strategies in order to prevent them. This allowed for the well-aimed optimization of the processing parameters resulting in a nearly void-free SLM processing of such maraging steel parts.Keywords: Selective Laser Melting, Direct Metal Laser Sintering, Lasercusing, Maraging Steel, heat balance, mechanical properties, Energy input, preheat temperature, void, defect1. IntroductionUnder the term additive manufacturing a variety of different technologies for generating parts from 3D-CAD-Data is combined.

IntroductionUnder the term additive manufacturing a variety of different technologies for generating parts from 3D-CAD-Data is combined. One of the most famous technologies of these is the selective laser melting process (SLM) which provides the layer wise production of high functional parts for testing applications and small series production (e.g. medicine and dental) out of metals. [GIB10], [HUA13], [KRU04], [ZAE11], [ZHA14]A special feature of the SLM-process and the additive manufacturing at itself is the high degree of geometrical freedom delivered and the opportunity of the process to produce parts without any kind of tools. These two factors, on the one hand the geometrical freedom, which can be used to apply a high grade of individualization in terms of for example personalization or weight reduction and on the other hand the opportunity to conduct production in flexible lot sizes without the need of tools.

Cause of these advantages additive manufacturing nowadays is used in both prototyping and small-series production. [CAM12], [GRZ11], [PET11]In the development of the SLM process there have been several material research topics (e.g. for aluminium, stainless steel or titanium) targeting an increase of the parts density and the void prevention. The Reason is that these factors affect the mechanical properties, such as strength and hardness, positively. It has been shown that parts processed via SLM can attain comparablemechanical properties to corresponding conventionally processed ones. [BUC11], [KON11], [KRU10], [ZHA14]The concrete influence of voids, respectively defects has been investigated in particular for the material Ti-6Al-4V. It has been found that the defects has a significant influence on both the static mechanical properties and the high cycle fatigue.

It has been found that the defects has a significant influence on both the static mechanical properties and the high cycle fatigue. [GRE13], [WCY14]The influence of various process parameters, such as energy input, on the porosity and meltpool appearance has been shown for both SLM and EBM process in comparison. It is stated that the energy input has a significant influence on the defect generation. For marginal parameters of too low and too high energy input the results are spattered defects on the lower side and round defects on the higher side. [GON13], [GON14], [STO14]In many cases the use of the geometrical freedom provided is limited due to material effects, such as high residual stresses and cracking, which result from high temperature gradients while processing the material in selective laser melting. [CAM12], [GRZ11], [FUL14]Resulting from this there is a variety of materials (e.g.

[CAM12], [GRZ11], [FUL14]Resulting from this there is a variety of materials (e.g. aluminium) hard to process without generating defects, such as pores and cracks, when not using a suitable preheating. Other strategies to obtain the high temperature gradients are to use scanning strategies different from the generally used stripe scanning. [BUC11], [KRU14]Additionally there were researches done, which have shown, that the preheating temperature has a significant impact on the resulting mechanical properties and the microstructure. As already stated above an additionally high impact on the mechanical properties is made by the scanning strategies, which result in different energy input levels and so different temperature gradients while processing. [BUC11]2.

[BUC11]2. ObjectiveDue to the lack of knowledge about the exact relationship between process parameters and the appearance of voids within the selective laser melting process it is basically difficult to determine which type of voids tolerable in terms of given requirements.The target of this study in this context is to find correlations of the process parameters energy input and preheat temperature, which are the main parameters defining the in-process heat balance, on the one hand and the resulting void appearance on the other hand for the given material Marage 300 (1.2709). Based on the findings the study furthermore should show the influence of the identified voids on the static mechanical properties and the resulting fracture surfaces.3. Experimental approach and methods3.1 MaterialFor the purpose of this study the material processed by SLM is Marage 300 (1.2709/AMS- 6514).

Experimental approach and methods3.1 MaterialFor the purpose of this study the material processed by SLM is Marage 300 (1.2709/AMS- 6514). Before processing an initial characterization of the powder has been done and shown below in Detail. The concrete investigations, which have been done are:Chemical analysis based on induced coupled plasma optical emission spectrometry(ICP-OES) (cp. figure 1) Fraction characterization based on Camsizer XT (Co. Retsch) using dynamic image analysis (cp. figure 2) Image acquisition based on scanning electron microscopy using SEM type Supra (Co. Zeiss) (cp. figure 3)Figure 1: Chemical analysis Marage 300 (ICP-OES)Figure 2: Fraction characterization Marage 300 (Camsizer XT)Figure 3: SEM-image of Marage 300 powderBased on the initial investigations shown in Figures 1-3 the powder used in the studies is completely characterized in terms of chemical composition, fraction and shape.

It fulfills on the one hand the chemical composition targets coming from AMS6514 (cp. figure 1) and shows theother hand a powder fraction of D10: 25,3 µm, D50: 40,8 µm and D90: 55,6 µm (cp. figure 2). Furthermore the powder shape is mostly spherical (cp. figure 3) with few adhesions.3.2 Machine and parametersThe machine used to process the specimen for the study is a Realizer SLM 100 with a 200 W Fiberlaser and Gaussian laser profile (general setup shown in Figure 4). Basically there are process parameters which has been fixed within the study (cp. figure 5) to gain the clear correlation between the variable parameters (cp.

Basically there are process parameters which has been fixed within the study (cp. figure 5) to gain the clear correlation between the variable parameters (cp. figure 6) stated.The variable parameters within the experiments are:preheat temperature, which is varied in 50 °C-steps between 25 °C and 300 °C - energy input, which is varied in 10 J/mm³-steps between 50 J/mm³ and 140 J/mm³The variation of the energy input, since it is no direct process parameter, is conducted by varying the scanning speed between 300 mm/s and 900 mm/s and using the following equation (1) to calculate the resulting energy input:𝐸𝑉 =𝑃𝐿 𝑣𝑆 ∗ ∆𝑦𝑆 ∗ 𝐷𝑆(1)Figure 4: General machine setupFigure 6: variable parametersFigure 5: Fixed parameters3.3 Analysis methodsThe specimen, which were used for the initial experiments are cubes of 10x10x10 mm. (cp. figure 7).

(cp. figure 7). In order to determine and correlate voids in microstructure the specimen produced by parameter variation the appearing voids has been identified by doing metallographic studies using light microscopy (cp. Figure 8).Figure 7: Overview specimen for initial studyAfterwards tensile testing has been done to correlate the appearing voids with their influence on the mechanical properties. The tensile testing has been done with in each case 5 specimen according to DIN 50125 Form A with the parameters 4x20 (cp. Figure 9) based on DIN ISO 6892-1 (cp. figure 10). Finally the tested tensile specimen were further investigated in scanning electron microscope (cp. figure 11) to identify the microstructural behavior and influence of the occurring voids while applying static load.Figure 8: Light microscope (Leica DM4000 M LED)Figure 9: Scheme of tensile specimen (acc.

figure 12) 4 different types of voids, which appear in different process windows, can be found. Basically they can be described by their shape the following:Spattered voids appear at a comparatively low energy input (50 J/mm³) and preheat temperature of 25 °CVirtually round voids appear at a comparatively high energy input (140 J/mm³) and preheat temperature of 25 °CCracks appear in very wide area of both energy input and preheat temperature.

Regarding the preheat temperature the area of appearance is 100 25 °C (for energy input of 95 J/mm³ to 105J/mm³) and 300 °C for 65 J/mm³ to 105 J/mm³The virtually void-free area appears between 65 J/mm³ and 85 J/mm³ at a preheat temperature of 25 °CThe parameters, which has been used for the further specimen building has been marked inred in figure 12.4.2 Tensile testingFigure 13: Results of tensile testingIn figure 13 the overview about the tensile testing of the specimen with the parameters chosen from the initial experiment is shown. It can be seen that there is a strongly variation of the different mechanical properties due to the voids appearing in the specimen.In detail it can be stated that the spattered voids has the biggest influence on the elongation to fraction, but in comparison to that the influence on the resulting strengths is very low.

The strength level of the specimen with spattered voids is the same as it is for the virtually void-free specimen, which gain the highest strength within the study. The specimen with virtually round voids show a significant decrease of both elongation to fracture and strengths.

It can be seen that the specimen has a fracture surface which is diagonal towards the axial orientation of the fracture. The specimen has nearly no deformation, but a high amount of spattered voids within, which are relatively homogeneous distributed over the whole specimen. The topography of the fracture surface is flat.Figure 15 shows the fracture surface investigated by electron scanning microscopy. On the left side an overview is given and shows a terraced and jagged fracture surface with many relatively big spattered voids. In the detail image on the right side the different appearing microfractographical characteristics. Marked with red circle and number 1 there is are free areas, which means an area which possibly has not been melted with the layers above, in different sizes. Area 2 shows some kind of a crack which goes parallel to the fracture surface.

Area 2 shows some kind of a crack which goes parallel to the fracture surface. Area 3 shows dimples, guides to ductile behavior of the material in this area.Figure 16: Microscopical image (Overview:left/Detail: right) of tensile specimen with parameter "virtually round voids"Figure 17: SEM-image (Overview:left/Detail: right) of tensile specimen with parameter "virtually round voids"In Figure 16 the longitudinal cut of one of the tensile specimens with virtually round voids is shown as an overview (left) and in detail (right). It can be seen that the specimen has a relatively flat fracture surface and is fractured perpendicular to its axial orientation. The topography of the fracture surface is jagged and the specimen has nearly little deformation, but contains a high amount of virtually round voids which are homogeneous distributed over the specimen.Figure 17 shows the fracture surface investigated by electron scanning microscopy.

It can be seen that the specimen is fractured perpendicular to the axial orientation of the fracture and the deformation is considerable compared to the “spattered void” and “virtually round void” specimen. The topography of the fracture surface is relatively flat with little areas which are kind of jagged. On the right image it can be seen that there are cracks within the fracture surface seem to run perpendicular to the fracture surface. The remaining area of the specimen is nearly void free, but for a big spattered void at around 2 mm below the fracture surface.Figure 19 shows the fracture surface investigated by electron scanning microscopy. On the left an overview is given and shows a relatively flat and only little jagged fracture surface with areas containing many linear voids and other areas being nearly void free. In the detail image on the right side the different appearing microfractographical characteristics.

In the detail image on the right side the different appearing microfractographical characteristics. In area 1 very small dimples can be seen, which means ductile areas of the material. Area 2 shows some kind of a secondary cracking which goes parallel to the fracture surface. Area 3 shows a free surface, which is possible not melted to the upper layer and additionally an enclosed cavity.Figure 20: Microscopical image (Overview:leftDetail: right) of tensile specimen with parameter "virtually void-free"Figure 21: SEM-image (Overview:left/Detail: right) of tensile specimen with parameter "virtually void-free"In Figure 20 the longitudinal cut of one of the tensile specimens which has been nearly void free is shown as an overview (left) and in detail (right). It can be seen that the specimen is fractured perpendicular to the axial orientation of the specimen.

It can be seen that the specimen is fractured perpendicular to the axial orientation of the specimen. Additionally it can be stated that the specimen has a relatively flat and only little jagged fracture surface with considerable deformation and a few spattered voids straight below the fracture surface. The rest of the specimen, as it can be seen on the right side is nearly void free.Figure 21 shows the fracture surface investigated by electron scanning microscopy. On the left side an overview is given and shows a flat and only little jagged fracture surface with in comparison to the prior specimen few mainly round and few spattered voids, which are concentrated in the middle of the fracture surface. The considerable deformation of the specimen results in a decrease of the investigated fracture surface. In the detail image on the right side the different appearing microfractographical characteristics are highlighted.

In the detail image on the right side the different appearing microfractographical characteristics are highlighted. Area 1 is showing free surfaces, which possibly guide to a layer which is unmolded with the above layer. Area 2 shows the already stated round to spattered voids appearing and number 3 an area of small dimples.Comparing the 4 different parameter setups resulting from the void identification of the initial experiments the results show that the exact processing parameters have a huge influence on the resulting density and appearing voids. On the other hand the further results from tensile testing and metallo- and fractography show that the different void types have a significant influence on both the mechanical properties and the appearance of the fracture surface after tensile testing.

In terms of mechanical properties spattered voids have the biggest influence on the elongation on fracture, which means a decrease of about 50% in comparison to the virtually void-free specimen. Virtually round voids and cracks in contrast have the biggest influence on the strengths. The decrease in strength is about 13% in case of tensile strength and 34% regarding the comparable yield strength of the virtually void-free specimen. Regarding the fracture surfaces the results show that there are voids in any of the specimen but that there appearance in terms of size and number varies from parameter setup to parameter setup. The spattered voids parameter setup for example is resulting in free surfaces, corresponding with unmolded neighbored layer, which are showing up in smaller fashion at the fracture surface of the virtually void-free parameter setup as well.

In contrast to that the virtually round voids parameter shows mainly unmolded powder particles and many round voids in the fracture surface. The cracks parameter setup in comparison shows cracks running parallel and perpendicular to the fracture surface as characteristics. All of the specimen has in common that there is always an area of dimples, which can possibly guide to ductile fracture behavior.5. ConclusionsThe characterization of the different kinds of the apparent voids for the maraging steel marage 300 subject to the energy input and the preheat temperature has been done. Therefore exact process windows for each void type has been identified and the influence of energy input and preheat temperature on the appearance has been shown. Regarding to that the identification and characterization of the apparent voids could been done.The correlation of the apparent voids and resulting mechanical properties from tensile testing has been done on parameter setup layer.

Furthermore the influence of the apparent voids in perpendicular building direction on the mechanical properties has been shown.Metallography and fractography has shown the different appearance of the voids in both longitudinal cut and fracture surface. The void types which has been identified through the initial experiment and following metallography has been found and confirmed in the metallo- and fractography of the tensile specimen. The parameter setup “virtually void-free” shows voids in the fracture surface, which guides to the conclusion that there is a further optimization and investigation necessary to determine the optimal processing parameters.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffects of Heat Treatment and Fast Neutron Irradiation on the Microstructure and Microhardness of Inconel 625 Fabricated via Laser-Powder Bed FusionT. Keya1, V. O'Donnell2, J. Lieben1, A. Romans1, G. Harvill1, M. Andurkar3, J. Gahl2, S.M. Thompson3, and B. C. Prorok11Department of Materials Engineering, Auburn University, Auburn, AL 36849 ²University of Missouri Research Reactor (MURR), University of Missouri, Columbia, MO 65211 3Alan Levin Department of Mechanical & Nuclear Engineering, Kansas State University, Manhattan, KS 66506AbstractThe microstructure of Inconel 625 fabricated via Laser-Powder Bed Fusion (L-PBF) was investigated in as-printed and heat-treated conditions.

The very high cooling rates inherent to the L-PBF process generally result in fine microstructures and complex residual stress fields which requires annealing to reduce stress and tailor the microstructure to obtain the desired mechanical properties. Inconel 625 alloy, a nickel-based superalloy, continues to be a common material employed with the L-PBF process. The unique microstructure produced by the L-PBF process and different phases introduced by different heat treatment processes require investigation to facilitate the material’s wide range of applications. This paper investigates the influence of heat treatments at 700°C, 900°C and 1050°C for one hour on the microstructure and microhardness of the L-PBF parts. The parts were irradiated using ‘fast’ neutrons in University of Missouri Research Reactor Center (MURR). The microhardness before and after radiation are also compared.1.

The parts were irradiated using ‘fast’ neutrons in University of Missouri Research Reactor Center (MURR). The microhardness before and after radiation are also compared.1. IntroductionLaser Powder Bed Fusion (L-PBF) offers the ability to fabricate metal parts with highly- customized geometries, tailored structures and consolidated assemblies. This additive manufacturing (AM) method has several advantages of traditional manufacturing methods, including, fabrication-friendly environments promoting lean production by reducing material waste, fabrication of parts made from a broad range of materials including superalloys, and reduction of fabrication time and cost [1-2]. During the L-PBF process, a focused high-energy laser beam selectively melts a powder layer accompanied with rapid cooling/solidification for a predetermined shape defined via computer aided drawing (CAD) data.

This process is followed by the coating of a new powder layer that is melted selectively to the previous layer. This cycle goes on until the final product is completely manufactured. During L-PBF, very high heat fluxes are imposed for creation of a melt pool which results in very high thermal gradients within the heat affected zone [3]. These thermal gradients manifest into significant residual stress distributions. During solidification, cooling rates can be as high as 103 K/s to 106 K/s, whereas for traditional methods, e.g. casting, has a cooling rate on the order of 1 to 103 K/s [4].

During solidification, cooling rates can be as high as 103 K/s to 106 K/s, whereas for traditional methods, e.g. casting, has a cooling rate on the order of 1 to 103 K/s [4]. These complex thermo- mechanical phenomena lead to fine dendritic inhomogeneous microstructures that require1036investigation; often to determine how an L-PBF material performs relative to its wrought counterparts.Inconel 625 (Alloy 625, IN625) is a nickel superalloy with exceptionally superior mechanical properties that derives its strength from solid solution strengthening by chromium (Cr), molybdenum (Mo) and niobium (Nb) atoms [4, 5]. In addition to solid solution strengthening, IN625 exhibits precipitation hardening primarily from the precipitation of the fine metastable γ″ phase.

In addition to solid solution strengthening, IN625 exhibits precipitation hardening primarily from the precipitation of the fine metastable γ″ phase. It is suitable for high temperature applications and is corrosion resistant making it an ideal candidate for aerospace, medical or energy sectors.Researchers have reported formation of γ″, δ, Pt2Mo-type phase, Laves phase and carbides in the typical FCC γ matrix of IN625 [6]. Floreen et al. reported formation of γ″ precipitates at 593℃-760℃ and formation of Laves and δ precipitates at 704℃-982℃ in wrought IN625 [6]. The stable δ phase, which can be needle or plate shaped, is formed upon dissolution of metastable γʺ precipitate due to local rises in Nb content [7,8]. Dinda et al. studied the microstructure of IN625 fabricated via Laser aided Directed Energy Deposition (L-DED) and reported a stable dendritic structure, even at 1000℃ [9]. Hu et al. found reduced ductility with increasing temperature in L- PBF IN625 [10].

Hu et al. found reduced ductility with increasing temperature in L- PBF IN625 [10]. This decrease in ductility is caused by intergranular cracking due to inhomogeneous microstructure introduced by the L-PBF process. Stoudt et al. reported faster precipitation in L-PBF IN625 compared to wrought material [11]. The nature of the L-PBF process causes localized concentration of Nb and Mo atoms that acts as the driving force to form δ precipitates at 870℃ in as quickly as 15 minutes. δ phase formation can be controlled by a stress relieving heat treatment and modifying Nb and Mo content. Another study on forged, L-PBF and Laser Melting Deposition (LMD) IN625 revealed heterogeneous microstructure with columnar dendrites in the latter two [12]. The microstructure in L-PBF IN625 is finer and more textured than that found in the forged and LMD microstructures.

The microstructure in L-PBF IN625 is finer and more textured than that found in the forged and LMD microstructures. The L-PBF microstructure can contribute to a material’s superior hardness and strength, as well as its poor ductility relative to wrought material. Heat treatment of IN625 parts at 1100℃ makes the mechanical properties more comparable to that of wrought parts. The superior mechanical properties found in L-PBF IN625 may also be attributed to high dislocation density along with the fine dendritic microstructure [13]. Marchese et al. reported formation of γʺ precipitates and Cr-rich carbides after heat treating L-PBF IN625 at 700℃ for 24 hours which caused high strength with poor ductility. Recrystallization occurred after annealing the material at 1150℃ for 2h resulting in dissolution of precipitates and a mixture of large as well as fine grains. Li et al. also reported high dislocation density as well as high residual stress in L-PBF IN625 [14].

Li et al. also reported high dislocation density as well as high residual stress in L-PBF IN625 [14]. In addition, the precipitation of the δ phase at 800℃ and 900℃ was found to result in higher hardness [14]. Precipitates or carbides usually are not found in as-built L- PBF IN625 [15]. Decrease in lattice constant after heat treatment was also found which is attributed to significant carbide precipitation.A detailed investigation of effects of nuclear radiation on microstructure and mechanical properties of AM IN625 can facilitate use of this alloy in nuclear reactors. Related studies found embrittlement and increase in hardness due to irradiation induced defects and precipitation. Hashimoto et al. conducted a study on ion irradiated Inconel 718 that was traditionally manufactured, solution annealed (SA) and precipitation hardened (PH) [16]. Samples were irradiated using Fe+, He+ and H+ beams of different strength at 200℃.

Samples were irradiated using Fe+, He+ and H+ beams of different strength at 200℃. Nanoindentation on irradiated samples revealed an increase in hardness for both SA and PH conditions. Higher dose and He injection on SA samples induced larger concentration of Frank loops and cavities resulting in irradiation-induced hardening. In Fe irradiated PH IN718, softening occurred due to incoherent γʹ and γʺ precipitates. In He irradiated PH samples, hardening was observed with increasing He1037concentration for up to 14 at%, then softening occurred at 20 at% He concentration due to dissolution of precipitates. Another study on Nimonic PE-16 (another nickel-based superalloy) alloy also emphasizes on He embrittlement over γʹ precipitation hardening [17]. Angeliu et al. suggested reduction in helium concentration and solute segregation to control radiation hardening [18]. Song et al.

Angeliu et al. suggested reduction in helium concentration and solute segregation to control radiation hardening [18]. Song et al. presented a comprehensive study on proton irradiated stainless steel, Ni based alloys and ferritic alloys [19]. Most samples showed formation of voids, dislocation loops and precipitates upon irradiation at 360℃ and 2.5 dpa. IN625 experienced enrichment of Cr and Mo on grain boundaries and precipitation of Pt2Mo-type phase induced by the irradiation. Ni-based alloys found to be more prone to radiation damage compared to ferritic alloys. Cieslik et al. reported traditionally manufactured IN625 to be more resistant to ion irradiation compared to its L-PBF counterpart [20]. Softening occurred for wrought alloy with increase in fluence whereas L- PBF IN625 experienced softening at 1.0 dpa and hardening at 3.0 dpa. The nano-hardness of the as-built L-PBF IN625 was found to be about 8% higher than that of wrought IN625 before irradiation.

The nano-hardness of the as-built L-PBF IN625 was found to be about 8% higher than that of wrought IN625 before irradiation. Increase in hardness for wrought IN625 was about 3%, but there was 6% increase in hardness for L-PBF IN625 after irradiation at 0.1 dpa. Softening at 1.0 dpa can possibly be attributed to grain growth and migration of defects such as voids, loops, and dislocations. Fincher et al. suggested post-radiation annealing to alleviate some radiation damage in solution annealed IN718 [21]. In-situ annealing at 300℃-500℃ for short period of time can relieve embrittlement induced by H+ and Ni2+ ions to some extent.The objective of this research work is to investigate the effects of different heat treatment schedules comprising of 700oC, 900oC, 1050oC for 1 hour on the microstructure and micro- hardness of IN625 fabricated by L-PBF. Metallurgical characterization is carried out on as-printed and heat-treated samples to analyze formation of precipitates.

Metallurgical characterization is carried out on as-printed and heat-treated samples to analyze formation of precipitates. Fast neutron irradiation was conducted on another similarly processed set of samples and micro-hardness was measured afterwards. Finally, sample hardness data were compared before and after irradiation. There is lack of studies that focus on microstructure and mechanical properties of additively manufactured IN625 irradiated by fast neutrons. This study will provide a better understanding of microstructural evolution of L-PBF IN625 before and after heat treatment and open doors to further investigate the effects of fast neutron irradiation on as-printed and heat-treated L-PBF IN625 for facilitating their use in low nuclear attenuation applications.2. Experimental Methods2.1 Material and L-PBF Process ParametersA Concept Laser MLab Cusing 100R L-PBF AM systems was used to produce all IN625 specimens.

Experimental Methods2.1 Material and L-PBF Process ParametersA Concept Laser MLab Cusing 100R L-PBF AM systems was used to produce all IN625 specimens. The AM machine was equipped with a 100 W fiber laser that has a wavelength of 1070 nm. The build volume of the machine measures 90 x 90 x 80 mm. Cubes with dimensions of 10x10x10 mm3 were additively manufactured. The laser power was set to 90 W with a scan speed of 800 mm/s. The laser diameter was 80 µm and the layer thickness is set to 25 µm with a hatch spacing of 60 μm. A density of 99.79% was achieved with these parameters.

The laser diameter was 80 µm and the layer thickness is set to 25 µm with a hatch spacing of 60 μm. A density of 99.79% was achieved with these parameters. The compositions of the Inconel 625 powder used for L-PBF and the wrought Inconel 625 are indicated in Table 1.1038Table 1: Elemental composition of Inconel 625 powder (% in weight)Element Ni Cr Mo FeNb+Ta Co CPSAlTi Mn% wtBal 219<5.0 3.5<1.0 < 0.1 <0.015 <0.015 <0.4 <0.4 <0.52.2 Heat TreatmentAfter separating the specimens from the built plate, each of them was exposed to a different heat treatment except one that was kept in the as-printed condition. Heat treatments were performed at 700℃, 900℃ and 1050℃ for 1h each. Temperature inside the furnace was measured using a K-type thermocouple and all samples were air cooled.Heat treatment parameters were chosen for precipitation of the metastable γʺ and stable δ phase. Fig. 1(a) shows the complete TTT diagram for wrought IN625 [22] and Fig.

Fig. 1(a) shows the complete TTT diagram for wrought IN625 [22] and Fig. 1(b) shows the thermal exposure conditions for δ formation in additively manufactured IN625 compared to wrought IN625 [11]. It may be seen from Fig. 1(b) that the time for δ phase formation for the same temperature is drastically reduced for the AM-LPBF samples compared to wrought. This behavior can be explained by the segregation of solute elements during solidification, which increases the probability for secondary phase formation. During L-PBF, the cooling rates are much higher than those experienced during conventional processes. This is known to lead to a finer dendritic structure with smaller dendrite spacing. Zhang et al.

During L-PBF, the cooling rates are much higher than those experienced during conventional processes. This is known to lead to a finer dendritic structure with smaller dendrite spacing. Zhang et al. showed that L-PBF produces a finer microstructure for IN625 than other AM processes (Electron Beam Melting and L-DED) and that the cellular/columnar substructures are enriched in Mo and Nb, while Ni and Cr are deficient [23].(a)(b)Figure 1: (a) Time-Temperature-Transformation (TTT) diagram of IN625, adapted from [22], (b) TTT diagram comparing the presence of δ-phase in additively manufactured IN625 to the wrought material. The dashed lines estimate approximately 1 %-volume fraction of δ-phase, adapted from [11]2.3 Metallographic Sample PreparationSpecimens were sectioned parallel to the build direction and hot mounted on CONDUCTOMOUNTTM mounting resin. Grinding and polishing were carried out on each sample1039down to 0.25 μm colloidal silica suspension.

Grinding and polishing were carried out on each sample1039down to 0.25 μm colloidal silica suspension. To reveal microstructure, samples were electro- etched with 70% phosphoric acid at 3-5 volts for 5-40 seconds.2.4 Characterization TechniquesScanning Electron Microscopy (SEM) was used to characterize microstructure before and after heat treatment on L-PBF specimens. A JEOL 7000F SEM unit was used to take high magnification images of the microstructure of the samples. Vickers hardness was also measured using a Phase-II Micro Vickers Hardness Tester before and after heat treatment and fast neutron irradiation.2.5 IrradiationA 25-week irradiation study was conducted on polished L-PBF samples in the as-printed and heat-treated conditions. ‘Fast’ neutrons produced from a cyclotron (GE PETtrace) at University of Missouri Research Reactor Center (MURR). This experiment looked at hardness value changes in the samples as the irradiations progressed.

This experiment looked at hardness value changes in the samples as the irradiations progressed. Samples were periodically removed for hardness testing over the course of the experiment to track variations. Samples were placed in a cyclotron vault, next to a fluorine-18 production target. The neutron flux inside the cyclotron vault had been modeled, showing that the neutrons resulting from the 18O(p, n)18F reaction have energies of 0.4 MeV to 16.5 MeV, and that the neutron flux in the vault is, conservatively, between 1.8 ×109 and 3.0 ×109 n cm-2 s-1 [24]. To standardize the results, the neutron flux was converted to 1 MeV equivalent neutrons, resulting in an equivalent flux between 2.4×109 and 4.0×109 n cm-2 s- 1 of mono-energetic neutrons.3. Results and Discussion3.1 Microstructure: As-printed conditionMicrostructures of the as-printed L-PBF IN625 specimens are shown in Fig. 2.

Results and Discussion3.1 Microstructure: As-printed conditionMicrostructures of the as-printed L-PBF IN625 specimens are shown in Fig. 2. Typical ‘U’ or ‘C’ shaped melt pools related to the laser scanning process are visible on the surface parallel to build direction. Although it is difficult to measure the exact width due to overlapping, the width of the melt pool is about 100-120 μm. The overlapping is also a result of the employed hatch spacing, thus resulting in relatively high density with no gaps between melt pools. This morphology is formed when the material rapidly solidify after melting by the laser source as soon as the laser moves. Li et al. also reported similar morphology that is more ‘V’ shaped resembling welding tracks [15]. The angle of the ‘V’ is a function of laser scanning speed and any increase in scanning speed decreases the angle and the melt pool becomes narrower.At higher magnification, cellular and columnar grains can be observed as shown in Fig. 2(b).

2(b). This may indicate growth of dendrites in different directions. An overall columnar growth along build direction is found as indicated by arrows on Fig. 2(b). The growth of the dendritic structure through melt pools is shown in Fig. 2(c). The dendrites are about 500 nm wide with Nb and Mo segregation on an interdentritic region and a Ni- and Cr-rich dendritic core as reported by other studies [14, 23].1040(a)(b)(c)Figure 2: SEM micrographs of as-printed L-PBF IN625 after electro-etching, (a) Melt pools on the plane parallel to build direction, (b) Cellular and dendritic microstructure at higher magnification, (c) Dendritic growth through melt pool boundaries at higher magnification.3.2 Microstructure: Heat Treated at 700℃Microstructure of the L-PBF specimens after 1 hour heat treatment at 700⁰C are shown in Fig. 3. At lower magnification, melt pools are noticeable after 1h (Fig. 3a) and it looks very similar to the as-printed microstructure (Fig. 2a).

3. At lower magnification, melt pools are noticeable after 1h (Fig. 3a) and it looks very similar to the as-printed microstructure (Fig. 2a). Dendritic microstructure and elemental segregation at the interdendritic region are more evident at higher magnification (Fig. 3b-c). These observations agree with other studies performed L-PBF IN625 [11, 13, 15]. Several research groups reported formation of metastable γʺ precipitate in the range of 550-750℃ [6-7, 25-28]. γʺ precipitate has a tetragonal structure with composition of Ni3(Nb, Ti, Al). These precipitates can be plate, disc, lens, spherical or elliptical shaped. It can take a long time to form γʺ phase in traditionally manufactured IN625, but precipitation may occur as soon as in 1h in AM materials [15]. Some areas observed in fig. 3(a, b) appear to be deficient of precipitation which may be attributed to elemental inhomogeneity throughout the sample.

Some areas observed in fig. 3(a, b) appear to be deficient of precipitation which may be attributed to elemental inhomogeneity throughout the sample. Areas with Nb depletion may not be able to form γʺ precipitates as carbides like NbC forms [6]. Precipitation is also affected by Ti content in the material.Build DirectionFigure 3: SEM micrographs of L-PBF IN625 after heat treating at 700℃ for 1 hour; (a) Melt pools on the surface parallel to build direction, (b) Melt pools with cellular and columnar growth at higher magnification, (c) Precipitation at interdendritic region.3.3 Microstructure: Heat Treated at 900℃Fig. 4 presents the microstructure after heat treatment at 900⁰C for 1h. Remnants of the melt pools can still be observed after 1h of heat treatment (Fig. 4a, b).

4 presents the microstructure after heat treatment at 900⁰C for 1h. Remnants of the melt pools can still be observed after 1h of heat treatment (Fig. 4a, b). While some precipitation is evident in certain areas, some areas have been recrystallized due to the dissolution of elemental segregation/precipitation that exist in the as-printed and 700℃ heat treated conditions. Plate and needle shaped particles are formed along sub-grain boundaries after 1h that resemble δ phase as shown in Fig. 4c [6, 29]. δ precipitates have orthorhombic structure with composition of Ni3Nb or1041Ni3(Nb, Mo) [6, 25]. It can be acicular, plate or globular shaped, although acicular, thin rod or needle shaped particles are most common. It is easier to detect and separate δ particles from its unique morphology. Laves phase forms at around the same temperature range in IN625, but it has a “blocky” irregular shape and its formation in IN625 is limited due to lower Fe content compared to IN718 [6].

Laves phase forms at around the same temperature range in IN625, but it has a “blocky” irregular shape and its formation in IN625 is limited due to lower Fe content compared to IN718 [6]. Stable δ phase forms from metastable γʺ precipitates when it dissolves into the matrix leaving high local concentration of Nb atoms [15, 29]. It usually starts forming at grain boundaries, then spreads out throughout the grain. Researchers have reported formation of δ phase formation at 650-982℃ [6, 25, 28, 29]. As can be seen from the TTT diagrams in Fig. 1, longer time might be required if the temperature is relatively low to form the precipitates. Baldan et al. reported evident formation of δ phase at 725℃ after 500h for traditionally manufactured IN625 [29]. Also shown in fig. 1, δ precipitates form much faster in AM materials compared to wrought IN625.

reported evident formation of δ phase at 725℃ after 500h for traditionally manufactured IN625 [29]. Also shown in fig. 1, δ precipitates form much faster in AM materials compared to wrought IN625. Due to the nature of the L-PBF process, there is Nb enrichment along interdendritic regions that triggers formation of precipitates that are Cr deficient, but enriched with larger Nb and Mo atoms [4]. Stoudt et al. reported formation of δ particles in L-PBF IN625 at 870℃ after 15 minutes [11]. Lass et al. reported formation of plate shaped δ precipitates at 870℃ for 1h in L-PBF IN625 [4].Figure 4: SEM micrographs of L-PBF IN625 after heat treating at 900℃ for 1 hour; (a) Melt pools on the plane parallel to build direction, (b) Columnar microstructure at higher magnification, (c) Precipitation in columnar regions.3.4 Microstructure: Heat Treated at 1050℃Microstructure of heat-treated sample at 1050℃ for 1h is presented in Fig. 5.

5. Melt pool boundaries, dendrites, elemental segregation, and most precipitates disappear at this heat treatment temperature causing complete recrystallization. Equiaxed grains with some annealing twins are visible. Presence of annealing twins indicates release of some residual stress that formed due to rapid cooling during the L-PBF process [12, 14]. Although not investigated in this study, some irregular block shaped MC, M6C or other carbides might be present after this heat treatment temperature [13, 15, 25, 26]. These carbides are often randomly distributed and have no specific orientation relative to the γ matrix [7]. Melt pools, precipitates and dendrites disappear usually at temperatures above 1000℃ [6, 12, 15].

Melt pools, precipitates and dendrites disappear usually at temperatures above 1000℃ [6, 12, 15]. Sufficient recrystallization with some primary carbides is reported to occur at 1150℃ in L-PBF IN625 [13,14, 15, 23].1042Figure 5: SEM micrographs of L-PBF IN625 after heat treatment at 1050℃ for 1h.3.5 Hardness: Before and After IrradiationVickers hardness data for the as-printed and heat treated samples before and after irradiation are shown in Fig. 6. It may be seen that the pre-irradiation hardness of the as-printed L- PBF sample is about 340.50±6.06 HV which is close to literature values [12, 15]. It is much higher than wrought IN625 which is about 260±5 HV [12].

It is much higher than wrought IN625 which is about 260±5 HV [12]. The high cooling rates during the L-PBF process cause “trapping” of Nb and Mo atoms and this phenomenon builds up residual stress in the material resulting in higher hardness compared to wrought or forged IN625 [12, 14, 15].Heat treating at 700℃ for 1h increases pre-irradiation hardness by about 5% which indirectly indicates γʺ precipitation [12]. Hardness before irradiation keeps decreasing with increases in temperature. Although δ precipitate causes hardening of material [15], the hardness decreases at 900℃. The effect of γʺ dissolution, release of residual stress and increase of grain size may dominate over that of δ phase formation [12, 14]. The hardness decreases at 1050℃ and becomes 266.6±5.02 HV which is close to wrought IN625 values found in the literature [12].

The hardness decreases at 1050℃ and becomes 266.6±5.02 HV which is close to wrought IN625 values found in the literature [12]. This phenomenon is caused by dissolution of precipitates and increase in grain size [12, 14, 15].As-printed and heat-treated L-PBF samples were irradiated using fast neutrons produced at MURR. Over the course of the experiment, hardness values of the samples were measured intermittently, as summarized in Fig. 6. The post-irradiation hardness of all samples display both softening and hardening, but they are relatively consistent after a fluence of 2.05×1015 n/cm2. All samples saw notable changes between the initial pre-irradiation measurements and the measurement taken after 8.57×1014 n/cm2 fluence. After that, hardness values did not show much change.

All samples saw notable changes between the initial pre-irradiation measurements and the measurement taken after 8.57×1014 n/cm2 fluence. After that, hardness values did not show much change. Relevant studies are not available that investigate effect of fast neutron irradiation on the microstructure and mechanical properties of L-PBF IN625, but some have reported irradiation hardening due to formation and rearrangement of defects and precipitates in the material [19, 20, 30, 31]. The presence of both hardening and softening effects is not uncommon [31]. Irradiation can give rise to local strain causing formation of loops that promote hardening [30]. It can also promote formation and growth of precipitates [19]. Defects like voids, loops or interstitials can form and accumulate resulting in increased hardness.

It can also promote formation and growth of precipitates [19]. Defects like voids, loops or interstitials can form and accumulate resulting in increased hardness. On the other hand, dissolution of precipitates, increase in grain size, or resettling of the defects can cause softening [20].The 1-MeV-equivalent neutron fluence at the end of the 25 weeks was 9×1015 n/cm2. This fluence would be considered low for a typical material damage study but was enough to get1043preliminary results that show this method of fast neutron irradiation is viable for future experiments.Figure 6: Vickers hardness as function of accumulated neutron fluence for L-PBF specimens.4. ConclusionsThe microstructure and microhardness of L-PBF IN625 before and after heat treatment have been investigated in this study. The microhardness after fast neutron irradiation is also presented and compared. Some important observations from this study are outlined below:1.

The microhardness after fast neutron irradiation is also presented and compared. Some important observations from this study are outlined below:1. L-PBF produces fine dendritic microstructure in IN625 that grows along the build direction due to rapid cooling.2. Melt pool boundaries and interdendritic segregation are retained through heat treatment at 700℃. At 900℃, melt pools start to dissolve and intergranular precipitation can be observed. 3. Melt pools, elemental segregation and precipitates are dissolved at 1050℃ and equiaxed grains with annealing twins are formed.4. As-printed L-PBF IN625 has higher hardness compared to wrought IN625. 5. L-PBF samples heat treated at 700℃ for 1h displayed the highest hardness among all samples before irradiation, whereas 1050℃ sample showed the lowest hardness among L-PBF samples and most comparable to wrought IN625.10446.

Both hardening and softening occur in L-PBF IN625 during continuous exposure to fast neutrons up to 9×1015 n/cm² normalized fluence.L-PBF technology can be a very powerful tool for modern nuclear reactor applications, but extensive research is required to investigate the performance of materials like IN625 with a complex microstructure produced using the technology. This study sheds light on evolution of microstructure of L-PBF produced IN625 superalloy in as-printed and heat-treated conditions at various temperatures which can be very useful to explain its mechanical properties like microhardness. Hardness after irradiation is also investigated after different levels of fluence which is a step forward towards understanding the performance of L-PBF IN625 in low nuclear attenuation applications.AcknowledgementThis material is based upon work supported by the U.S.

Department of Energy’s Office of Nuclear Energy under Award Number DE-NE0008865.DisclaimerThis report was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights. Reference herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof. The views and opinions of authors expressed herein do not necessarily state or reflect those of the United States Government or any agency thereof.References1.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEffects of Identical Parts on a Common Build Plate on the Modal Analysis of SLM Created MetalTristan Cullom1, Troy Hartwig2, Ben Brown2, Kevin Johnson3, Jason Blough3, Andrew Barnard3, Robert Landers1 Douglas Bristow1, Edward Kinzel11Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409. 2Kansas City National Security Campus, Kansas City, MO 64147. 3Department of Mechanical Engineering – Engineering Mechanics, Michigan Technological University, Houghton, MI 49931.ABSTRACT The frequency response of parts created with Additive Manufacturing (AM) is a function of not only process parameters, powder quality, but also the geometry of the part.

Modal analysis has the potential to evaluate parts by measuring the frequency response which are a function of the material response as well as the geometry. A Frequency Response Function (FRF) serves as a fingerprint of the part which can be validated against the FRF of a destructively tested part. A practical scenario encountered in Selective Laser Melting (SLM) involves multiple parts on a common build plate. Coupling between parts influences the FRF of the parts including shifting the resonant frequencies of individual parts in ways that would correspond to changes in the material response or geometry. This paper investigates the influence of the build plate properties on the coupling phenomena. This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1.

INTRODUCTION Recent improvements in AM have allowed various materials to be used to create specimens as well as creating complex geometries. With AM increasing in usage, small-scale manufacturing is becoming less prominent, resulting in the need for validation techniques. There are NDE methods in place, such as Computed Tomography, which are used to detect porosity or other geometry defects; however, they are expensive [1]. A cost effective means that is used to detect defects is modal analysis and it has been applied to various structures in the automotive, civil infrastructure, and aerospace industries. [2] Modal analysis can be used to validate parts on the basis of their resonant frequencies. If a part is less dense than another, the resonant frequency will shift upward due to the inverse relationship between the resonant frequency and the density of the part.

If a part is less dense than another, the resonant frequency will shift upward due to the inverse relationship between the resonant frequency and the density of the part. A concern arises however, when multiple parts are tested modally on a common build plate that is used during the SLM process and the result is a coupling effect that presents itself similar to a tuned mass damper. To show this effect a simulation was conducted in which one rectangular cantilever was in the middle of the build plate and another where two were symmetric about the vertical axis and were separated by one inch. Figure 1 shows the resulting FRFs of the single and two cantilever systems.

Figure 1 shows the resulting FRFs of the single and two cantilever systems. To showcase the sensitivity of the frequency response, the same two geometries were used but instead the stiffness of the parts was decreased by one percent.2254Figure 1: Single versus two cantilever frequency response.From the graph, the mode shapes of the two cantilever system includes a mode shape where the two parts are out of phase from each other in addition to an in phase shape. A problem that can be seen in validating parts is that if a part is defective in terms of stiffness, the resulting FRF of a single cantilever is approximately the same in terms of frequency as the nominal two cantilever system.

Consequently, validating multiple parts becomes troublesome since it is almost impossible to distinguish between a two nominal cantilevers and a single defective cantilever.A basic way to address this issue is to print each part separately and run the modal test; however, time becomes a major problem since it will obviously increase the testing time. Another common way to approach this problem is to ensure that the parts that are printed on the same build plate have natural frequencies that are far away from each other. The problem with this approach is that it eliminates the possibility of “fingerprinting” or comparing parts against nominal parts as a means of removing bad parts from a batch.This study serves the purpose of identifying the underlying reasoning behind the coupling effect and investigates ways to improve upon current modal analysis techniques as a means to test AM parts.

Simulation tests were designed and analyzed to determine the relationship between the coupling-response as a function of the underlying frequency of the build plate first mode as well as the deformation change of the build plate as a function of its material properties.22552. SIMULATION SET-UP AND RESULTS To explore the coupling effect, the first mode of the build plate was found using ANSYS and can be seen in Figure 2. The geometry of the build plate used was 250 mm × 250 mm × 12.7 mm and modeled as 1010 steel. The first mode shows clear node and anti-node lines which will be utilized in tests that will be explained later in this report with the theory that the coupling response between cantilevers is dependent on their location on the build plate itself.XYFigure 2: Build plate first mode from ANSYS simulation.The first natural frequency of the build plate itself was 671.66 Hz.

The cantilever geometry was designed to have the first natural frequency of the cantilever resonate at the same frequency as the build plate. The geometry of the cantilevers was 0.375” x 0.15” x 2.61”. The properties of the cantilevers were set to be that of as-printed 304L stainless steel, with a young’s modulus of 1.8 x 1011 Pa, a density of 7.85 g/cm3, and a Poisson’s ratio of 0.24. The two cantilevers were centered about the Y axis and were positioned at 1/3 and 2/3 along the X axis.The height of the cantilevers was varied from 95% to 102.5% of the nominal length. The first three eigenfrequencies were recorded which represented the two coupled mode shapes as well as the build plate first mode. Figure 4 shows the eigenfrequencies of the first three mode shapes as a function of the length ratio of the current height to the nominal height.The first eigenfrequency was the build plate and the next two were the in phase and out of phase modes of the two cantilevers.

This behavior remained true until the heights of the cantilevers reached 98.5% of the nominal length; the eigenfrequencies started transitioning and the coupled responses between the cantilevers reached its maximum. When the cantilevers were at the nominal length, the order of modes was now the coupled response between the two cantilevers followed by the build plate mode. Figure 3 shows a surface plot of the eigenfrequency response from the cantilevers as well as the eigenfrequencies imposed on top of the surface plot.

Figure 3 shows a surface plot of the eigenfrequency response from the cantilevers as well as the eigenfrequencies imposed on top of the surface plot. As expected, the frequency response of the system had the same shape as the eigenfrequency plot.2256Figure 3: Surface plot showing the velocity frequency response with eigenfrequency black lines imposed above surface plot.To see how the response of the structure changes as the cantilevers approached the build plate’s resonant frequency, five FRFs were recorded at critical ratios: 0.95, 0.9675, 0.985, 1 and 1.025 Lo, seen as the vertical black lines in the surface plot.

Figure 4 through Figure 8 show the respective FRFs of each of the five length ratios.Figure 4: L/Lo = 0.95 Frequency Response2257Figure 5: L/Lo = 0.9675 Frequency ResponseFigure 6: L/Lo = 0.985 Frequency Response2258Figure 7: L/Lo = 1 Frequency ResponseFigure 8: L/Lo = 1.025 Frequency ResponseLooking at the first graph, the two modes of the cantilevers are approximately 130 Hz above the build plate, resulting in little interaction between the modes. As the length of the cantilevers continued to increase, the build plate frequency shifted and interacted the most at 98.5% the nominal length, seen in how the magnitudes in the first and third mode being equal.

Similarly, as the length continued past that of the nominal length, the build plate frequency shifted above that of the cantilevers.To determine the relationship between the properties of the build plate and the respective coupling between the cantilevers and the build plate, the properties of the build plate were varied for the next series of tests. Knowing that the resonant frequency is proportional to the square root of the ratio of young’s modulus to the density, Equation 1, three configurations of cantilevers were used, seen in Figure 10, and the resulting eigenfrequencies of the structure were recorded.2259fE ∝ ρ= (a)= (b)= (c)Figure 9: Cantilever Configurations.These three configurations encompass the highest and lowest locations of deformation in the build plate: the corners and middle, respectively. The motivation for this is that the coupling should be highest when the difference of deformation is highest.

The motivation for this is that the coupling should be highest when the difference of deformation is highest. Referring to Equation 1, the two main tests that were conducted were keeping the ratio of stiffness to density to be the same but varying their magnitudes by the same amount, and varying the stiffness while keeping density to be nominal. For each test the difference between configurations (b) and (c) against (a) were recorded as a function of the build plate stiffness as well as the deformation of the build plate.The first test involved varying the stiffness to density ratio from 0.1 to 100 times the nominal ratio. With this test the first modal frequency of the build plate should theoretically stay the same whereas the cantilever frequencies should change due to the change in deformation of the build plate.

With this test the first modal frequency of the build plate should theoretically stay the same whereas the cantilever frequencies should change due to the change in deformation of the build plate. Figures 10 and 11 show the difference between the configurations when the stiffness to density ratio was constant.In addition to plotting the eigenfrequency difference, a theoretical difference of density was plotted as well considering studies have already down that a difference in engineering properties because of processing parameters also yields a difference in eigenfrequencies.

[4]2260(1)Figure 9: Eigenfrequency difference between configurations (b) and (a) along with theoretical density difference as function of stiffness ratio.Figure 10: Eigenfrequency difference between configurations (c) and (a) along with theoretical density difference as function of stiffness ratio.From the graphs, as the build plate’s stiffness and density both increased, the difference between the configurations converged towards zero. This is due to both the build plate frequency shifting above that of the cantilevers as well as the deformation of the build plate decreasing as the stiffness of the plate increased.The next test involved varying the stiffness from 0.5 to 100 times the nominal stiffness but having the density remain constant at the nominal value. Figures 12 and 13 show the eigenfrequency difference as well as the theoretical density difference.By varying only the stiffness, the build plate frequency varied below and above the cantilever frequencies.

This creates two significant differences from the first test: the build plate and cantilevers couple together as well as the mode shape of the build plate being affected as the stiffness increases. The change in the mode shape is the reasoning behind why even though the2261build plate frequency was increased above that of the cantilevers, the difference between the configurations does not approach zero as the stiffness continuously increases.Figure 11: Eigenfrequency difference between configurations (b) and (a) along with theoretical density difference as function of stiffness ratio.Figure 12: Eigenfrequency difference between configurations (c) and (a) along with theoretical density difference as function stiffness ratio.The second test results are heavily dependent on the ratio of the stiffness.

As stated earlier, the mode shape of the structure as the stiffness increases changes to the point to where the difference between the frequencies crosses effectively zero and then converges towards a constant value as the stiffness continues to increase. From this, the ability to differentiate between two cantilevers and a single cantilever is difficult if the build plate is stiffened without adding appreciable mass.The results from the third test do not show similar behavior as the first test: the coupling response did not converge toward 0 as the build plate frequency became farther and farther away from that of the cantilevers. In addition, the only points at which the difference in eigenfrequencies was small was due to the mode shape change which is unmerited due to the transition point being at the discretion of the processing the simulation results since there is no clear ratio that signifies2262when the mode shape changes.

Some reasons as to why the build plate frequency trend does not remain true is to the fact that the build plate’s mass approaches that of the cantilevers when its frequency increases; this causes the cantilevers and build plate to not be independent of each other in terms of mass. Because of this, the influence of the cantilevers shifts the frequency of the build plate even more; smearing the eigenfrequency results.With the three tests in mind, a potential solution to the coupling effect between parts can be minimized by increasing the build plate frequency above that of the cantilevers by increasing both the stiffness and density. An obvious way to achieve this solution is to simply make the build plate thicker. To test the effectiveness of increasing the build plate thickness, the eigenfrequencies of configurations (a) and (c) were recorded and plotted as a function of the additional build plate thickness.

To test the effectiveness of increasing the build plate thickness, the eigenfrequencies of configurations (a) and (c) were recorded and plotted as a function of the additional build plate thickness. Figure 14 shows the resulting eigenfrequencies of both configurations.Figure 13: Eigenfrequencies as function of build plate thickness.The difference between the first and second eigenfrequencies of configuration (c) against configuration (a) was inversely proportional to the square root of the thickness of the build plate. With this in mind, tolerances of 0.1% and 0.05% between frequencies were chosen to show the behavior of the configurations as the thickness of the build plate changed; the resulting thicknesses that yielded these success criterions were 2.75” and 7.625”, respectively.

Figure 15 shows the FRFs of the configurations at nominal, 2.75”, and 7.625” additional thickness.2263Figure 14: FRFs of configurations (a) and (c) as thickness of build plate changed.There are two distinct trends that occur in Figure 15. The first being that as the build plate thickness increased, the out of phase amplitude became insignificant to the point of not being seen in the FRF. The second trend being that the difference between configurations progressed from 0.84 to 0.2 to 0.1 Hz. With this in mind, the difference between the eigenfrequencies can be reduced, making it easier to fingerprint defective from nominal parts.With the results from increasing the geometry of the build plate itself, it begs the question about the impact of bolting the plate to a heavier object. A plate was created that was 60” × 60” and attached to the bottom of the nominal build plate.

A plate was created that was 60” × 60” and attached to the bottom of the nominal build plate. The thickness was set to be 1” and the frequency response for both configurations was recorded and can be seen in Figure 15.Plate ModeCantilever Mode(s)Figure 15: FRF of (a) and (c) with plate attached to bottom of build plate.2264As predicted, the coupling effect was reduced to the point to where the FRFs of (a) and (c) were almost on top of each other. The underlying reasoning for this can be seen in the deformation of the build plate itself. Since the build plate was bolted to a heavier mass, the deformation of the build plate was reduced substantially. With bolting the build plate to a heavier plate, the simulation results almost match the same shape as the theoretical boundary condition results. The only difference is that with the 60” by 60” plate, it had a resonant frequency near that of the cantilevers, resulting in the second peak seen in both configurations.

The only difference is that with the 60” by 60” plate, it had a resonant frequency near that of the cantilevers, resulting in the second peak seen in both configurations. The only downside to this behavior is that the driving force required to displace the stiffening plate would require powerful equipment. Further research will be conducted to optimize the topology of the stiffening plate to encompass lattices to minimize the overall weight of the plate, while keeping the stiffness to density ratio relatively constant to maintain the same frequencies of the cantilevers attached to the build plate.To showcase the scenario of the deformation from a boundary condition perspective, the build plate was modeled with a fixed support on the top and bottom surface. Figure 16 shows the resulting FRFs of (a) and (c).

Figure 16 shows the resulting FRFs of (a) and (c). As predicted, with the deformation of the build plate set equal to zero, the difference between the resonant frequencies is also zero.Figure 16: FRF of configurations (c) and (a) with fixed boundary condition.22653. SUMMARY AND CONCLUSIONSFrom the findings in this report it is certain that increasing the stiffness of the common build plate that parts are attached to mitigates the coupling effect. To facilitate this behavior, the thickness of the build plate was varied as well as bolting the build plate to a heavier plate. The results from both of these simulations show that the coupling behavior was minimized with bolting the build plate to a heavier plate as opposed to changing the geometry of the build plate itself. As a direct result, validating parts in terms of frequency response is made more accurate by increasing the stiffness of the build plate by meaning of bolting it to another structure.4.

As a direct result, validating parts in terms of frequency response is made more accurate by increasing the stiffness of the build plate by meaning of bolting it to another structure.4. ACKNOWLEDGMENTSThis work was funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes. This work contains a patent pending technique, U.S. Patent Application Serial No. 14/941,258.5. REFERENCES[1] Frazier, W. E., “Metal additive manufacturing: a review.” Journal of Materials Engineeringand Performance, 23(6), 1917- 1928, 2014.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffects of In-Situ Mechanical and Chemical Polishing on Surface Topography of Additively Manufactured Fiber-Reinforced Polymers Aman Nigam, Bruce L. Tai Department of Mechanical Engineering Texas A&M University, College Station, TX 77843, USAAbstractAdditive manufacturing of fiber-reinforced polymers (FRPs) has revolutionized fused filament fabrication (FFF) by producing polymeric parts with enhanced mechanical properties. However, FFF suffers from poor surface quality and dimensional accuracy, particularly for FRPs, due to their abrasive and rheological nature. This examines an in-situ polishing scheme for FRPs in the FFF configuration. Glass-fiber-reinforced Nylon was used as the study material.

This examines an in-situ polishing scheme for FRPs in the FFF configuration. Glass-fiber-reinforced Nylon was used as the study material. Three polishing schemes, mechanical, chemical, and a combined thereof, were adopted along with various parameters in each case. The results show significant surface improvements in all cases, and the combined process can further reduce the Ra value to around 2 μm and the dimensional error to 0.2 mm and less. The combined process also enhances surface uniformity (i.e., similar Ra in all directions). In particular, with the combined approach, the in-situ polishing scheme is expected to improve the quality of 3D printed FRPs significantly.Keywords: Fiber-reinforced polymer, Fused Filament Fabrication (FFF), Hybrid process, Polishing, 3D Printing, Surface roughness.1. IntroductionIn recent times, polymer and polymer composites have drawn significant attention in manufacturing owing to their excellent mechanical, physical, and thermal properties.

IntroductionIn recent times, polymer and polymer composites have drawn significant attention in manufacturing owing to their excellent mechanical, physical, and thermal properties. Fiber- reinforced polymers (FRPs) are characterized by properties such as high strength to weight ratio, high modulus, high fracture toughness, corrosion, and thermal resistance [1]. Due to these advantageous properties, FRPs have become a cost-effective alternative to traditional metals in aerospace, marine, oil and gas, and process industries [2]. Hand lay-up is the most popular and widely used open-mold, fiber composite manufacturing technique. However, the production rates are low, and both material and labor costs are high [3]. Further, finishing processes that involve machining and trimming are often required and essential to meet the design requirements [4].

Further, finishing processes that involve machining and trimming are often required and essential to meet the design requirements [4]. As the application field of FRPs expands, the opportunity for improving and modifying the current manufacturing techniques has increased for its fabrication.The last decade witnessed an exponential development in additive manufacturing, and it continues to grow due to its economic and versatile nature. The use of FRPs has transformed additive manufacturing into a robust technique and has enabled the prototyping of highly customized parts with enhanced properties. The solid model is sliced into individual layers in this automated manufacturing method, which gets printed in a stacked fashion. FRPs can be additively manufactured by fused filament fabrication (FFF) with either short (chopped) or continuous fiber configurations.

FRPs can be additively manufactured by fused filament fabrication (FFF) with either short (chopped) or continuous fiber configurations. Although continuous fiber configuration offers excellent mechanical properties in1661certain directions, the major drawbacks include the lack of design flexibility and supplier-defined cloud-based slicing software [5]. In comparison, short-fiber printing remains the most popular and versatile approach in the additive manufacturing of FRPs.The surface quality of FFF manufactured parts is typically poor and depends on the layer resolution and material properties [6]. This drawback of FFF, when coupled with the abrasive nature of FRPs, leads to inferior surface topography [7]. Conventional machining methods such as milling are typically used to enhance the surface finish of these materials.

Conventional machining methods such as milling are typically used to enhance the surface finish of these materials. However, due to the abrasive and anisotropic nature of FRPs, machining of such materials often fail by matrix cracking, fiber pull-out, swelling, and delamination [8]. Kuram studied the micro-machinability of polyamide 6 (Nylon), and glass-fiber reinforced polyamide 6 manufactured by injection molding. Tool wear, cutting forces, and surface roughness were evaluated at different spindle speeds and feed rates. It was concluded that forces increased with feed rate and spindle speed, but surface roughness decreased with spindle speed for both workpiece materials. It was also pointed out that the friction between the tool and the workpiece material can induce heat generation at the contact area, leading to the softening of the polymer, thus impairing surface quality [9]. Further, Parandoush et al.

Further, Parandoush et al. examined the mechanical finishing process for 3D printed carbon fiber reinforced polymers composites via CNC milling. The surface roughness of original and finished samples was compared at various cutting depths. A 70% reduction in Ra was reported with a smooth, consistent, and predictable surface morphology after finishing [10]. These studies were successful to a certain extent; however, the post-processing requires multiple machines to fabricate and polish, making them expensive, time-consuming, and less flexible. Also, manual finishing of these FRPs results in health-related issues due to the harmful dust formation [11,12].In addition to improving the surface finish by mechanical means, extensive work has also been published in the chemical polishing of 3D printed polymers, but not specifically for FRPs.

Most of the studies were carried out either by immersing the printed part in a suitable chemical solvent or introducing it in a vapor environment in hot or cold conditions [13,14]. The solvent erodes the surface's polymer material, and the surface tension forces drive a reduction in the surface roughness. Kuo et al. developed a device to chemically polish ABS printed parts using acetone vapors. The surface finish improvement was found to be more than 88%, and similar results were observed for parts with different curvatures [15]. Nigam et al. presented a novel concept of in-process mechanical and chemical polishing of fused filament fabrication (FFF) and examined its feasibility in improving surface finish and dimensional accuracy. More than an 85% reduction in surface roughness was observed with significantly reduced form errors for ABS [16].

More than an 85% reduction in surface roughness was observed with significantly reduced form errors for ABS [16]. While the results from these investigations are vital in reducing the surface roughness, they are primarily limited to simple polymers such as ABS. No or very limited work has been reported for more functional, fiber-reinforced polymer composites.Henceforth, considering the limitations of past research and challenges of finishing additively manufactured FPRs, this study aims to examine the effects of mechanical, chemical, and combined polishing of FRPs in an in-situ additive manufacturing environment. A custom three-axis hybrid manufacturing setup was built to provide precise motion control and repeatability of polishing parameters.

A custom three-axis hybrid manufacturing setup was built to provide precise motion control and repeatability of polishing parameters. A glass fiber-reinforced polymer (GFRP) in short fiber configuration was1662used, and a factorial analysis was conducted to quantify the effects of spindle speed, feed rate, chemical polishing depth, and speed on surface roughness and dimensional accuracy. Further, statistical analysis was conducted to determine the optimum parameters in individual polishing schemes to design a combined mechanical plus chemical polishing approach.To structure this paper, the experimental setup, methods, and analysis techniques are explained in Section 2. Section 3 enlists and discusses the results obtained after experiments for multiple cases. Finally, section 4 draws the concluding remarks and the future scope of this research.2.

Section 3 enlists and discusses the results obtained after experiments for multiple cases. Finally, section 4 draws the concluding remarks and the future scope of this research.2. Experimental Setup and Testing Methodology2.1 Setup and Test SpecimenTo fulfill the objectives of this research, it was essential to conduct polishing in a precise and repeatable manner. Experiments were carried out using the setup shown in Fig.1. A hybrid manufacturing setup with simultaneous printing and polishing capabilities was built, integrated with a servo-assisted motion control system (Moog Animatics, USA). A high-speed spindle (NSK, Japan) with a carbide end mill (3.175 mm diameter, three flutes) was employed for mechanical polishing at different spindle speeds. Additionally, a pen-style device filled with a suitable chemical solvent (formic acid for nylon-6) was used for chemical polishing.

Additionally, a pen-style device filled with a suitable chemical solvent (formic acid for nylon-6) was used for chemical polishing. This device was connected to a reservoir with a flow control valve to regulate the rate of chemical dispensing to ensure a controlled erosion process. This method controls both the solvent dispensing rate and the normal force applied on the surface. More details are discussed in the subsequent sections.Figure 1: Custom-built setup with high precision 3-axis motion control stage, high-speed spindle for mechanical polishing, and a chemical polishing felt assembly.1663For this study, specimens were printed in 30% (mass fraction) short glass fiber reinforced nylon-6 composite (Owens Corning, USA) using the FFF technique. The printing temperature was 245oC, and the heated platform temperature was 80oC. Manufacturing was done in an acrylic enclosure to maintain uniform temperature and prevent material warping.

The printing temperature was 245oC, and the heated platform temperature was 80oC. Manufacturing was done in an acrylic enclosure to maintain uniform temperature and prevent material warping. Cuboidal and cylindrical test specimens were produced to test polishing in straight-walled and curved surfaces. The dimensions are listed in Table 1. To make a direct comparison, the first 7 mm of the sample was left unpolished.Table 1. Dimensional parameters for different specimen geometries. Cylinder 20 mm 12 mm 0.2 mm 5 mm (25 layers) 5Parameter Final desired X-Y dimension Total height Printed layer height Polished height Number of layers polished in 1 passCuboid 20 mm 12 mm 0.2 mm 5 mm (25 layers) 52.2 Polishing Control StrategyIndividual and combined mechanical and chemical polishing were performed utilizing an automated approach. Custom G-codes (Simply3D, USA) were generated by varying the polishing parameters according to the design of experiment.

Custom G-codes (Simply3D, USA) were generated by varying the polishing parameters according to the design of experiment. Cuboidal and cylindrical samples were polished along their outer periphery while the top and bottom surfaces were left as printed.To automate the process, the part is manufactured using a two-step scheme [16]. This process employs a pseudo part at a distance from the primary printed part. The pseudo part is the outer shell of the primary print with modified dimensions. As shown in Fig. 2 (a), the tool offsets (distance between nozzle and the respective polishing tools) are applied to polish the primary part when the pseudo part is being printed. After a designated set of layers is polished, the printing continues on top of the partially polished primary part, and the cycle repeats itself.As shown in Fig.

After a designated set of layers is polished, the printing continues on top of the partially polished primary part, and the cycle repeats itself.As shown in Fig. 2 (b), to reach the desired final dimension compensating for the tool radius and the depth of polishing, the primary part is first printed at a larger dimension and then polished. For example, in the case of mechanical polishing for a desired final dimension of 20 mm, the primary part was designed to be 21 mm (i.e., 20 mm plus 2 two depths of cut of 0.5 mm) while the pseudo part was 23.18 mm (i.e., 20 mm plus a tool diameter). Similarly, for chemical polishing, the primary part was designed to be 20.2 mm (for a maximum pressing depth = 0.1 mm), while the pseudo part was 24.95 mm with a tool diameter of 4.95 mm.1664Figure 2.

In-situ polishing strategy (a) polishing of primary part when pseudo part is being printed; (b) Tool radius and radial depth of cut compensation in the design process.2.3 Parameters for Design of ExperimentIn mechanical polishing, the cutting parameters should be meticulously selected as they directly affect the cutting forces, affecting the surface roughness. The chip load is the theoretical length of material fed into each cutting edge as it moves through the work material. As the chip load increases, it requires more force to shear the material being cut. Suppose the cutting forces were high while machining, the printed part adhered to the print platform may get disturbed, which may be detrimental to the surface finish. The chip load, therefore, was maintained below 0.05 mm. Also, to hold the workpiece material on the heated platform, down milling was performed.

The chip load, therefore, was maintained below 0.05 mm. Also, to hold the workpiece material on the heated platform, down milling was performed. In down milling, cutting forces are directed downwards, thus pressing the workpiece rather than lifting it.The design of experiment for mechanical polishing, as shown in Table 2, consisted of two levels of each spindle speed and the feed rate. The dependent variable is the surface roughness of the produced part.Table 2. Design of experiment for mechanical polishing.Case 1Case 2Case 3Case 4Spindle Speed (rpm)20000200003000030000Feed Rate (mm/min)80012008001200Chip load (mm)0.0130.020.0090.013In chemical polishing, the material erosion rate depends on the concentration of the chemical solvent and the deposition rate. The chemical deposition rate is controlled by the chemical flow rate and the speed of polishing.

The chemical deposition rate is controlled by the chemical flow rate and the speed of polishing. In this experiment, the flow rate was kept constant while the speed of polishing was varied. Moreover, the radial force that the polishing felt applies on the printed part is defined by the pressing depth (directly proportional relation) [17]. Two pressing depths and polishing speeds were selected for comparison, as shown in Table 3.1665Table 3. Design of experiment for chemical polishing.Case 5Case 6Case 7Case 8Pressing Depth (mm)0.050.050.100.10Feed Rate (mm/min)600800600800In addition, optimum parameters from individual mechanical and chemical polishing were used to design a combined polishing strategy to investigate its performance compared to the individual polishing techniques. In this approach, mechanical polishing was performed first, which was followed by chemical polishing.

In this approach, mechanical polishing was performed first, which was followed by chemical polishing. Mechanical polishing acts as bulk material removal process and hence preceded chemical polishing, which can be characterized as finishing.2.4 Surface Roughness Characterization and Optical MeasurementsThe manufactured samples were examined in macro and micro scales. The macro-scale analysis was conducted to assess the dimensional accuracy using Mitutoyo PH A41-profile projector (Kanagawa, Japan). The surface topography was analyzed for micro-scale analysis using Bruker ALICONA Infinite Vision (Raaba, Austria) on an area of 2×2 mm2 using a 0.8 mm Gaussian filter (ISO-4288-1996). The Ra value was measured along and perpendicular to the feed direction for all polished faces. Both methods are based on non-contact measurement, suitable for polymer composites.3.

The Ra value was measured along and perpendicular to the feed direction for all polished faces. Both methods are based on non-contact measurement, suitable for polymer composites.3. Result and DiscussionSurface measurement results for mechanical, chemical, and combined polishing arepresented in the following three sections.3.1 Surfaces after Mechanical PolishingThe surface topography of the GFRP composites manufactured using FFF 3D printing is much different from the topography of the ones made through hand layup. The surface texture arises not only because of the abrasive nature of the material but also from the stair-stepping effect attributed to the layered printing. As shown in Fig. 3, the unpolished region of the part displays noticeable layering lines with fibers protruding out of the outer periphery. The average surface roughness (Ra) was measured across the four faces against the layering lines of each sample.

The average surface roughness (Ra) was measured across the four faces against the layering lines of each sample. It was computed to be 22.52 m for cuboid and 27.27 m for cylindrical samples.1666Figure 3. Mechanical polishing (a) cuboidal sample with unpolished and polished faces; (b) microscopic image of the polished face; (c) surface topography of the polished face; (d) microscopic image of the unpolished face; (e) surface topography of the unpolished face.Generally, during milling of polymer composites, the fibers are exposed to shear stresses, inducing fiber failure at the cutting surface in the form of brittle fracture across the fiber, fiber pull-out and fiber-matrix debonding (by tensile fracture). This results in an overall smooth finish. After mechanical polishing at a depth of cut of 0.5 mm, the layer lines were eliminated, demonstrating a visibly smooth surface across all samples.

This results in an overall smooth finish. After mechanical polishing at a depth of cut of 0.5 mm, the layer lines were eliminated, demonstrating a visibly smooth surface across all samples. Compared to the unpolished faces, mechanically polished faces in cuboid samples had sharper corners. However, both cuboidal and cylindrical specimens displayed exposed fibers protruding out of the machined surface, as shown in Fig. 3. This is also a common phenomenon in FPRs machining.Fig. 4 shows the effect of polishing parameters on the average surface roughness (Ra) for cuboid and cylindrical samples. To judge the impact of polishing, Ra was measured alongside feed and perpendicular to the feed direction (across the layer lines). A total of four measurements were made for each sample across four different faces.

A total of four measurements were made for each sample across four different faces. Further, analysis is discussed below:1667Figure 4: Mechanical polishing- variation of Ra with spindle speed and feed rate; (a) Ra measured along feed direction; (b) Ra measured perpendicular to feed direction.Paired T-test reveals no significant difference in the surface roughness of cuboidal and cylindrical samples for the same polishing parameters (p-value = 0.57 at 95% significance level). This indicates that mechanical polishing provides similar results irrespective of the geometry of the workpiece sample.After mechanical polishing, a significant difference exists in the Ra measured alongside feed and perpendicular to the feed direction for both the geometries (p-value cuboid = 0.01, p-value cylinder = 0.03 at 95% significance level).

The surface roughness measured perpendicular to the feed direction is higher, suggesting that mechanical polishing is more effective in the feed direction.Further from Fig. 4, it can be inferred that the value of Ra increases with the feed rate and decreases with the spindle speed irrespective of the measurement direction. The best results were obtained for the polishing parameters of Case 3, which corresponded to the lowest chip load. Case 3 showed an 85.3% reduction for cuboid and an 85.45% reduction for cylindrical geometries compared to the unpolished region.Despite the smallest chip load (0.009 mm) with a theoretical roughness of 0.129 μm, the actual roughness is in twice the orders of magnitude larger. This indicates the limitations of roughness reduction by milling on FPRs.16683.2Surfaces after Chemical PolishingAs shown in the microscopic images of Fig. 5, the chemically polished face exhibits an overall smooth surface finish, and the layer lines were eliminated.

5, the chemically polished face exhibits an overall smooth surface finish, and the layer lines were eliminated. Unlike mechanical polishing, after chemical polishing, the fibers were encapsulated in the eroded polymer giving it a more uniform appearance. However, material accumulation in the form of surface beads was visible, which may negatively affect the surface finish.Figure 5. Chemical polishing (a) cuboidal sample with unpolished and polished faces; (b) microscopic image of the polished face; (c) surface topography of the polished face; (d) microscopic image of the unpolished face; (e) surface topography of the unpolished face.Fig. 6 shows the effect of polishing parameters on the average surface roughness (Ra) for cuboid and cylindrical samples. Polishing trials were conducted at pressing depths of 0.05 and 0.10 mm, and feed rates of 600 and 800 mm/min.

Polishing trials were conducted at pressing depths of 0.05 and 0.10 mm, and feed rates of 600 and 800 mm/min. Further, analysis is discussed below:1669Figure 6: Chemical Polishing- variation of Ra with pressing depths and polishing speed; (a) Ra measured along feed direction; (b) Ra measured perpendicular to feed direction.Paired T-test reveals no significant difference in the surface roughness of cuboidal and cylindrical samples for the same polishing parameters (p-value = 0.59 at 95% significance level). This indicates that chemical polishing provides similar results irrespective of the geometry of the workpiece sample.After chemical polishing, no significant difference exists in the Ra measured alongside feed and perpendicular to the feed direction for both the geometries (p-value- cuboid = 0.09, p-value- cylinder = 0.08 at 95% significance level). This implies that unlike mechanical, chemical polishing is non-directional and effective in both directions.Further from Fig.

This implies that unlike mechanical, chemical polishing is non-directional and effective in both directions.Further from Fig. 6, it can be inferred that minimum Ra was obtained for Case 8 (high pressing depth and high speed). Case 8 showed an 82.72% reduction for cuboid and an 86.24% reduction for cylindrical geometries compared to the unpolished region.On comparing the average results from Case 3 of mechanical polishing and Case 8 of chemical polishing, no significant difference exists in the Ra (p-value- cuboid = 0.98, p- value- cylinder = 0.27 at 95% significance level). This implies that chemical polishing reduces the Ra to a similar level to that of mechanical polishing at optimum parameters. However, it should be noted that their surface topographies are different.16703.3 Combined Mechanical and Chemical PolishingFrom the previous sections on individual mechanical and chemical polishing, it can be observed that they have some limitations.

Mechanical polishing is directional, while chemical polishing leads to the formation of surface beads. Therefore, this section examines a combined polishing approach and compares it to the previous results.Optimum parameters obtained from the previous sections for individual mechanical and chemical polishing were used to design a combined polishing strategy to investigate its performance. The parts were manufactured using a three-step scheme- after printing a set of layers, mechanical polishing (spindle speed: 30000 rpm, feed rate: 800 mm/min) was conducted, which was followed by chemical polishing (pressing depth: 0.10 mm, feed rate: 800 mm/min).As shown in Fig. 7, one part of each geometry was manufactured using the combined polishing method. The Ra was measured alongside feed and perpendicular to the feed direction (across the layer lines). The protruded fibers after machining were encapsulated in the eroded polymer.

The Ra was measured alongside feed and perpendicular to the feed direction (across the layer lines). The protruded fibers after machining were encapsulated in the eroded polymer. These encapsulated fibers were also shorter than those in the chemical-polishing-only cases. Further, the surface appeared to have fewer surface beads, meaning a more uniform erosion on a machined surface. Further analysis is discussed below:Figure 7. Combined polishing (a) cuboidal sample; (b) microscopic image of the polished face; (c) surface topography of the polished face; (d) cylindrical sample; (e) microscopic image of the polished face; (f) surface topography of the polished face.1671Paired T-test reveals no significant difference in the surface roughness of cuboidal and cylindrical samples for the same polishing parameters (p-value = 0.14 at 95% significance level).

This indicates that the combined polishing technique provides similar results irrespective of the geometry of the workpiece sample.After combined polishing, no significant difference exists in the Ra measured alongside feed and perpendicular to the feed direction for both the geometries (p-value- cuboid = 0.07 p-value- cylinder = 0.10 at 95% significance level). This implies that combined polishing is non-directional and effective in both directions.Compared to the unpolished region, the cuboidal and cylindrical samples showed 89.8% and 90.1% reduction, respectively. The average Ra measured for cuboidal and cylindrical were 2.47 m and 2.93 m, respectively.As summarized in Fig. 8, combined polishing displayed maximum reduction in Ra irrespective of the measurement direction and part geometry.Figure 8.

8, combined polishing displayed maximum reduction in Ra irrespective of the measurement direction and part geometry.Figure 8. Summary- Effect of various polishing configurations on the surface roughness3.4 Dimensional AccuracyApart from the microscopic analysis, macroscopic analysis was also performed to assess the dimensional accuracy of the polishing process for the samples polished with the most optimum parameters. The changes to the length and the width of the samples were measured for the two geometries. The results are summarized in Table 4. The dimensional accuracy for all cases has less than 1% error from the desired size. In fact, most of these dimensions are more accurate than a typical FFF 3D printing machine (0.1-0.5 mm). This suggests that the proposed polishing method can retain or even improve the dimensional accuracy.1672Table 4.

This suggests that the proposed polishing method can retain or even improve the dimensional accuracy.1672Table 4. Measured dimensional error for different various polishing configurations.Designed DimensionsMechanical PolishingChemical Polishing CubeCubeCylinderCylinder Length 20 mm, Width 20 mmCombined Polishing Cylinder CubeMeasured Length (mm) Measured Width (mm)20.12 20.0720.09 20.0720.02 20.0320.04 20.1120.14 20.0120.14 20.01Length Error % Width Error %0.6 -0.350.45 -0.350.1 -0.150.2 -0.550.7 -0.050.7 -0.054. ConclusionThis parametric study examined an in-process polishing technique to enhance the surface finish of fiber-reinforced polymer parts manufactured through the FFF 3D printing technique. Experiments were conducted on a custom-developed 3-axis hybrid manufacturing setup with the capabilities of performing in-situ surface polishing.

Experiments were conducted on a custom-developed 3-axis hybrid manufacturing setup with the capabilities of performing in-situ surface polishing. First, individual mechanical and chemical polishing was conducted, and the average surface roughness was measured for different polishing parameters and geometries. In addition, optimum parameters from individual polishing techniques were selected to design a combined polishing approach.The results found that the employed polishing method plays a vital role in determining the average surface roughness. Both individual mechanical and chemical polishing leads to more than an 80% reduction in the Ra compared to the unpolished face. However, unlike chemical polishing, mechanically polished surfaces were found to be directional. Further, at optimum parameters derived from the individual polishing schemes, combined polishing outperformed them with more than a 90% reduction in Ra of the unpolished face.

Further, at optimum parameters derived from the individual polishing schemes, combined polishing outperformed them with more than a 90% reduction in Ra of the unpolished face. Also, no directionality was observed in the case of combined polishing.Despite the apparent advantages of in-process polishing on GFPRs, many technical challenges remain and determine the future scope of this research. First, this study was based on glass fibers; it will be interesting to investigate the polishing behavior of other fiber materials or delicate mesh geometries. In the case of delicate mesh geometries, while conducting chemical polishing, the lack of radial force can be substituted with an increased solvent flowrate achieving a similar level of material erosion. However, while conducting mechanical polishing, reducing the depth of cut and chip load (hence the cutting forces) might be advantageous. This will be similar to micro-milling.

However, while conducting mechanical polishing, reducing the depth of cut and chip load (hence the cutting forces) might be advantageous. This will be similar to micro-milling. Secondly, a higher magnification analysis is required to examine the fiber failure mechanism post polishing. Most importantly, it is critical to determine the effects of surface polishing on the mechanical performance of these GFRP manufactured parts.This conclusion leads to the necessity for future expansion and an in-depth investigationof the in-situ process on its material removal mechanism.1673References[1] Farshbaf Zinati, R., & Razfar, M. R. (2013). An investigation of the machinability of PA 6/nano-CaCO3 composite. The International Journal of Advanced Manufacturing Technology, 68(9-12), 2489–2497. https://doi.org/10.1007/s00170-013-4875-3[2] Palanikumar, K., Karunamoorthy, L., & Karthikeyan, R. (2006).

In the current study, a numerical model for predicting the temperature distribution and residual stress in Directed Energy Deposited (DED) Ti–6Al–4V parts is utilized for determining a relationship between local part temperature gradients with generated residual stress. Effects of time-interval between successive layer deposits, as well as layer deposition itself, on the temperature gradient vector for the first and each layer is investigated. The numerical model is validated using thermographic measurements of Ti-6Al-4V specimens fabricated via Laser Engineered Net Shaping® (LENS), a blown-powder/laser-based DED method. Results demonstrate the heterogeneity in the part’s spatiotemporal temperature field, and support the fact that as the part number, or single part size or geometry, vary, the resultant residual stress due to temperature gradients will be impacted.

As the time inter-layer time interval increases from 0 to 10 second, the temperature gradient magnitude in vicinity of the melt pool will increase slightly.IntroductionDirected energy deposition (DED) is an additive manufacturing (AM) technique for building parts via the layer-wise addition of material [1]. Material, in the form of wire or powder, is fed into the directed energy source, which may be either an electron beam or laser. For laser- based DED, or direct laser deposition (DLD), a common system for use is Laser Engineered Net ShapingTM (LENS). During a typical LENS process, metallic powder is injected into a relatively high power laser (typically Nd:YAG) via multiple nozzles (typically four); all occurring in an inert, argon atmosphere. Due to the small laser spot size and high laser power, injected metallic powder melts forming a molten pool, or melt pool.

Due to the small laser spot size and high laser power, injected metallic powder melts forming a molten pool, or melt pool. The melt pool solidifies near-instantaneously upon removal of laser irradiation, and extremely high initial cooling rates then occur. A CNC stage below the laser, which is mounted with a substrate, moves in a pre-defined path along one plane, allowing for the creation of a single layer. The stage is then incrementally dropped until a final part is generated. The localized heating and extreme cooling rates experienced during LENS will175influence the microstructure, mechanical properties and residual stress of final metallic parts [2]– [5].To investigate the influence of DED process parameters on the properties of metallic parts, many have resorted to experimental or numerical means.

Since the heat transfer and temperature fields are relatable to observed microstructures and residual stresses in as-built LENS parts, thermal characterization of the process is highly pertinent. For experimental characterization, some have employed substrate-embedded thermocouples [6], [7] or infrared thermography [8], [9]. Numerical characterization of the LENS process has been performed using finite element analysis (FEA) [10], [11]. Griffith et al. [6] experimentally demonstrated that dimensional accuracy in the horizontal build plane is better compared to accuracy in the vertical fabrication direction for LENS. Hofmeister et al. [8] used thermography and FEA to measure and calculate the temperature gradient in the part. Costa et al. [10] studied the effects of idle time and substrate size on properties of deposited material.

[8] used thermography and FEA to measure and calculate the temperature gradient in the part. Costa et al. [10] studied the effects of idle time and substrate size on properties of deposited material. If the idle were eliminated or the substrate size was chosen to be very small, the chance of remelting previous layers was found to increase, thus causing process instability and dimension inaccuracy.Residual stress in DED parts can be detrimental to their mechanical performance during application and manufacture (e.g. part breaks during fabrication) [12], [13]. In addition, it can impact the net shape of DED parts via bowing and bending [14]. Residual stress, although important for mechanical integrity of DED parts, is difficult to measure. However, one can predict residual stress formation in DED parts through thermo-mechanical FEA; by coupling the part temperature field (and heat transfer) with its mechanical strain [14]–[16].

However, one can predict residual stress formation in DED parts through thermo-mechanical FEA; by coupling the part temperature field (and heat transfer) with its mechanical strain [14]–[16]. Heigel et al [14] has shown that to predict the residual stress accurately, modeling forced convection is necessary.One can make inferences on residual stress locations/magnitude within the part by quantifying the temperature gradient [15], [17]. The temperature gradient is important as it indicates (1) magnitude/direction of local conduction heat transfer (i.e. Fourier’s law) and (2) possible regions of solid expansion/contraction in vicinity of melt pool. Thermal cycling of a particular solid region occurs while new, neighboring layers are deposited and/or thermally agitated, and such cycling can cause repetitious expansion/contraction of the material.

Temperature gradients are dependent on the employed laser scan strategy during DED, as the relative laser scan direction will influence heat flux directionality. Therefore, one can strategically select DED process parameters to better minimize residual stress formation in produced parts.The study herein provides insights into the relationship between the user-defined laser scan strategy and resultant temperature field/gradients in rectangular parts fabricated via DED. Additionally, effects of deposition of multiple layers on thermal response on the first layer has been studied. Finally, effects of adding time interval after the fabrication of each layer on thermal response in each layer has been investigated.Physical ModelThe laser-based/blown-powder DED, e.g. LENS, process is physically modeled herein. For numerical modeling, material deposition has been modeled using either the quiet or inactive methods [18].In quiet methods, elements are present.

LENS, process is physically modeled herein. For numerical modeling, material deposition has been modeled using either the quiet or inactive methods [18].In quiet methods, elements are present. The thermal conductivity and heat176capacitance assigned to such elements are chosen to be near zero. In this case, conduction in these regions is very limited. The advantage of this method is that it is very easy to implement. On the other hand, for the inactive method, elements are included as they are added to the part. This method increases the accuracy of the model, although it is harder to implement. In this study the inactive method is chosen in order to increase the accuracy of the model.In order to simulate the LENS process and subsequent temperature, Ti, of the as-processed part, the three-dimensional transient heat equation shown in Eq.

(1) was solved numerically for each ith continua (melt pool, solid part, substrate, deposited powder) comprising the part during manufacture over the time window t0 - 1, i.e.:𝐶i(Ti)|𝑡o−1𝜕Ti(𝑥, 𝑦, 𝑧, 𝑡) ∂𝑡|𝑡o= ∇𝑘i(Ti)|𝑡o−1∇Ti(𝑥, 𝑦, 𝑧, 𝑡)|𝑡owhere Ci is the heat capacity (J/kg∙°C) and ki is the thermal conductivity (W/m∙°C). The continua distribution/phase is coupled with the unknown temperature field; therefore, it is evaluated at previous time and updated based on local temperature. In this study, laser power is assumed to be Gaussian. The Gaussian distribution is formulated as [19]:𝑄𝑟 =2𝑃𝛼λ 2𝐻 𝜋𝑟0(1 −𝑧 𝐻) exp (1 − (𝑟 𝑟0)2)where Qr is the input energy density (W/mm3), P the laser beam power (W), αλ is the the hemispherical, spectral absorptance of the powder and melt pool, r0 is the radius of the laser beam, H is the maximum depth, r is the current radius and z is the current depth.

The incident laser power is compensated by absorption coefficient of the powder, heat loss due to convection and radiation, as well as any loses due to scattering in the powder cloud [20]. The remainder of energy is absorbed by the powder for melting, and is also utilized for sensibly heating the melt pool and previously deposited layers.During LENS, powder is injected into a region within and preceding the melt pool for subsequent melting. During this relatively fast heat transfer process, the powder is assumed to be packed flush with the deposited layer thickness.

During this relatively fast heat transfer process, the powder is assumed to be packed flush with the deposited layer thickness. The effective thermal conductivity of the powder, kpb, was then calculated for a given temperature using [21]:𝑘pb 𝑘g= (1 − √1 − 𝜙) (1 +𝜙𝑘r 𝑘g) + √1 − 𝜙 {21 −𝑘g 𝑘s[11 −𝑘g 𝑘sln (𝑘s 𝑘g) − 1] +𝑘r 𝑘g}where kg is the thermal conductivity of the entrapped gas, which is only argon in most cases, and ϕ is the porosity of the powder, i.e.:𝜙 =𝜌s − 𝜌p 𝜌s177(1)(2)(3)(4)where ρs and ρp are the density of parts and powder (kg/m3).

The effective thermal conductivity of the irradiated powder, kr, which depends on thermal radiation exchange between neighboring particles and the particle-to-particle distribution/size [21], was found using: 3 𝑑p 4𝜀𝜎T𝑖,𝑜 1 − 0.132𝜀3 𝑑p =𝑘r|T𝑖,𝑜 = 4𝐹0𝜎T𝑖,𝑜where Fo is the radiation view factor, σ is the Stefan-Boltzmann constant, dp is the mean powder diameter and ε is the constant emissivity for a typical, spherical powder arrangement [21]. The effective thermal conductivity of the melt pool, kmp, was estimated using:𝑘mp|T𝑖,𝑜= (1 − 𝜃)𝑘l,eff + 𝜃𝑘s|T𝑖,𝑜where θ is the solid-phase volume fraction during melt pool formation; varying between 0 and 1, i.e.:θ =Vs Vs + Vl=Vs VmpHeat transfer due to Marangoni and natural convection within the superheated melt pool [22], [23] was accounted for by using an effective, temperature-independent liquid thermal conductivity, kl,eff [24]. Effective thermal conductivity used for melt pool in this is 20 W/m∙°C [25].

Effective thermal conductivity used for melt pool in this is 20 W/m∙°C [25]. Evaporation heat transfer within the melt pool was neglected.Heat transfer via convection and thermal radiation, between continua and surroundings, was estimated by assuming the shielding gas and chamber walls were at a uniform, constant temperature, T∞, i.e.:𝑞𝑐𝑜𝑛𝑣,𝑛𝑒𝑡 = ∬ ℎ̅(Ti − 𝑇∞)𝑑𝐴 𝑞rad,net = 𝜎𝜀(Ti4(𝑥, 𝑦, −𝑚∆𝑙, 𝑡) − 𝑇∞4)where ℎ̅ is the heat transfer coefficient (W/m2∙°C), σ the Stefan-Boltzmann constant (W/m2∙°C4) and ε, the emissivity of the part. Heat transfer between the parts and chamber gas was idealized as forced convection describable via an average heat transfer coefficient, ℎ̅. Based on Heigel et al., the heat transfer coefficient was selected as 10 W/m2∙°C [14].Table 1.

Based on Heigel et al., the heat transfer coefficient was selected as 10 W/m2∙°C [14].Table 1. Temperature dependent thermal properties of solid Ti–6Al–4V [14]T (°C)2093205 250 315425500540650760800870k (W/m∙°C)6.67.39.19.710.6 12.6 13.9 14.6 17.5 17.5 17.5 17.5Cp (J/kg∙°C) 565 565 574 586 603649682699770858895959Since the LENS process consists of mass and energy transport occurring at relatively fine spatial and temporal scales, e.g. micrometers and microseconds, respectively, highly resolved178(5)(6)(7)(8)(9)discretization of governing energy equations is required. This can be a challenge due to computational resource availability and time restrictions. A 3072 core cluster setup was utilized to expedite the simulation process while maintaining a fair level of accuracy.

A 3072 core cluster setup was utilized to expedite the simulation process while maintaining a fair level of accuracy. Simulations were performed using the commercially available software, COMSOL Multiphysics® 5.1 in conjunction with custom coding scripts.A tetrahedral meshing scheme was employed for spatial discretization of the part. Meshes consisted of cells with length approximately ~33% that of the laser diameter and a volume of approximately 1,000,000 μm3. Time was discretized into equal time steps of Δt = 1 ms. The solution was verified as mesh independent and independent of time by doing a sensitivity study. The magnitude of the temperature gradient between neighboring nodes separated by x, y, z was estimated using a second-order central difference.VerificationThe accuracy of the developed numerical model was assessed by simulating the LENS processing of Ti-6Al-4V performed by Marshall et al. [26]; parameters are summarized in Table 2.

[26]; parameters are summarized in Table 2. In this experiment, temperature measurements of the melt pool were conducted using a dual wavelength (DW) pyrometer. The DW pyrometer possessed a 6.45-micron/pixel resolution and a 4-mm field of view and was set to capture images with an acquisition frequency of 6.4 Hz. The as- received DW pyrometer was calibrated using a high-temperature tungsten filament at known temperature [26]. An OPTOMEC LENSTM 750 machine with 1 kW Nd:YAG laser was utilized. The substrate was made from Ti–6Al–4V and its size was 153 × 153 × 3.3 mm3. During the experiment, a multi-layer, single track wall was fabricated. The length, height and thickness of the wall is provided as 47.81 mm, 27.56 mm and 1.78 mm, respectively. The temperature was measured at multiple locations along the wall at each layer. Note that the numerical model employed herein uses temperature dependent material parameters and accounts for phase change.Table 2.

Note that the numerical model employed herein uses temperature dependent material parameters and accounts for phase change.Table 2. LENS process parameters [26] used to validate current numerical modelMaterial Ti–6Al–4VLaser power Scan speed Laser spot size Powder feed rate290 W12.7 mm/s2 mm0.32 g/sThe predicted and experimentally-measured temperature at the middle of a layer, for the first five layers, is presented in Table 3. It is observed that the temperature prediction from the employed simulation is ≈ 50 °C higher relative to results from experiments. Several reasons could explain these differences. In the model, effects of heat transfer inside the melt pool, such as Marangoni effects, were neglected. In addition, local convection heat transfer effects were neglected.Table 3. Peak temperature of middle point on the thin wall from simulation and experiment conducted by Marshall et al.

In addition, local convection heat transfer effects were neglected.Table 3. Peak temperature of middle point on the thin wall from simulation and experiment conducted by Marshall et al. [26]Layer 3 1730 °C 1745 °C 1740 °C 1760 °C 1770 °C Simulation Experimental 1707 °C 1708 °C 1748 °C 1705 °C 1717 °C1245179Numerical results and discussionRectangular parts with 6 × 6 mm2 cross-sectional dimensions were simulated while employing a ‘unidirectional’ scan strategy is shown in Fig. 1. As shown, the scan direction rotates 90° clockwise for each new layer compared to previous layer. After extensive experiments, the set of process parameters chosen that shown to produce dense parts. The laser power is set to 350 W and travel speed is set at 16.9 mm/s. The feed rate is 133 mg/s. The hatching space and thickness of each layer are 0.508 mm. The measurement location is chosen to be at the middle of the part’s first layer. In Fig.

The feed rate is 133 mg/s. The hatching space and thickness of each layer are 0.508 mm. The measurement location is chosen to be at the middle of the part’s first layer. In Fig. 2, the temperature response during the fabrication of the second and third layer is shown. It can be seen that peak temperature is located at deposition point.Figure 1. Scan strategy for (a) first layer, (b) second layer, (c) third layer, (d) fourth layer.Figure 2. Temperature response during deposition of (a) second layer and (b) third layer.In Fig. 3, temperature response and temperature gradient of measurement location (middle of first layer) is shown. It is observed that temperature oscillation will decrease at measurement180location during deposition of top layers. The reason is that distance between heat source and measurement location increases.

The reason is that distance between heat source and measurement location increases. In addition, due to heat accumulation, at the beginning of deposition of each new layer, temperature of the measurement location is higher compared to beginning of deposition of previous layer. Finally, temperature gradient during deposition of first layer is highest.Figure 3. (a) Temperature response and (b) temperature gradient magnitude of measurement location during fabrication of first five layers while employing a time interval of ~ 0 s.In Fig. 4, the temperature response and temperature gradient during deposition of first three layer is shown. Measurement location is located at the center of each layer during the deposition. It can be seen that peak temperature is nearly the same in all three layers, despite the fact that the temperature of the part is rising. However, temperature gradient magnitude of Layer 3 is lower than Layer 1.Figure 4.

However, temperature gradient magnitude of Layer 3 is lower than Layer 1.Figure 4. (a) Temperature response and (b) temperature gradient of measurement locations during fabrication of first three layers while employing a time interval of ~ 0 s.181In Fig. 5, temperature gradient during the deposition of first layer along x, y and z direction at measurement location is shown. It concluded that peak temperature gradient is highest along z direction. In addition, temperature gradient along the laser direction is higher compared to temperature gradient along the direction perpendicular to the laser direction.Figure 5. Temperature gradient along (a) x direction, (b) y direction, and (c) z direction. during deposition of first layer at measurement location.Finally, in Fig. 6, temperature response and temperature gradient during deposition of the first five layers of measurement location is shown. In this simulation between deposition of each layer, the laser was idle for 10 s.

In this simulation between deposition of each layer, the laser was idle for 10 s. It could be seen that although the initial temperature at the beginning of deposition of each layer decreased, peak temperature has not changed. On the other hand, the temperature gradient has decreased.Figure 6. Temperature response and temperature gradient of measurement location during fabrication of first five layers while employing a time interval of 10 s.ConclusionsThe temperature response during fabrication of parts produced by the DED, and specifically the Laser Engineered Net Shaping (LENS), method has been investigated. The proposed182continuum model for DED heat transfer assumes that the powder possesses an effective thermal conductivity that accounts for convection and radiation effects between particles and entrapped gas. The model is essentially conduction-based. Effects of fluid motion inside the melt pool have not been considered, and pore formation was neglected.

The model is essentially conduction-based. Effects of fluid motion inside the melt pool have not been considered, and pore formation was neglected. The absolute temperature field has been used to allow inference on residual stress formation in DED parts. Based on the numerical simulation, the below conclusions can be made:1) Temperature of the part will rise during the deposition of each layer. 2) The magnitude of the temperature gradient at the center of each layer decreases as the layer number increases; away from the substrate.3) As the inter-layer time interval increases, the temperature gradient in vicinity of the melt pool will increase.These findings further show that the geometry/size/number of parts can impact their temperature response, i.e. thermal history, as well as their temperature gradients, during DED.AcknowledgementsThis material is based upon work supported by the National Science Foundation under Grant No. 1563423.

thermal history, as well as their temperature gradients, during DED.AcknowledgementsThis material is based upon work supported by the National Science Foundation under Grant No. 1563423. Authors acknowledge the support of Mississippi State University’s Center for Advanced Vehicular Systems (CAVS) and High Performance Computing Collaboratory (HPC²). This manuscript was prepared while Scott Thompson and Nima Shamsaei were Assistant Professors at Mississippi State University.References[1]J. Mazumder and H. Qi, “Fabrication of 3D components by laser-aided direct metal deposition,” in Proceedings of the SPIE., 2005, vol. 5706, pp. 38–59.[2] M. . Griffith, M. . Schlienger, L. . Harwell, M. . Oliver, M. . Baldwin, M. . Ensz, M. Essien, J. Brooks, C. . Robino, J. . Smugeresky, W. . Hofmeister, M. . Wert, and D. . Nelson, “Understanding thermal behavior in the LENS process,” Mater. Des., vol. 20, no. 2–3, pp. 107–113, 1999. F.-J. Kahlen and A.

FECTS OF LASER WINDOW DEGREDATION ON LASER POWER AND DISTRIBUTION IN LASER SINTERINGBen Fulcher, David K. LeighHarvest Technologies Belton, Tx 76513AbstractLaser power is a key parameter in the laser sintering (LS) process, and tight control on laser power is necessary to produce quality parts with desirable mechanical properties. Unfortunately, temperature limitations hinder real-time monitoring and feedback of laser power within the process chamber. Therefore, in order to maintain consistent laser power during an LS build, the laser window, which provides a barrier between the processing chamber and the laser housing, must remain clean throughout the build. However, material outgassing leads to the buildup of condensation on the window, thereby reducing the amount of energy imparted to the powder bed. The buildup of condensation also necessitates frequent cleaning of the laser window and significantly reduces its life.

The buildup of condensation also necessitates frequent cleaning of the laser window and significantly reduces its life. Thus, laser window replacement is a large source of cost in a production environment. To compensate for the loss of laser power through the window, current practice is to steadily increase the laser power at the laser source during the build. This practice can be largely inaccurate, as it is difficult to predict the loss of laser power through the window at various stages in a given LS build. Thus, to achieve consistent mechanical properties in this manner, a trial and error-based approach is used. The study presented in this paper aims to characterize laser power and distribution for various levels of laser window degradation. In addition, methods to reduce or eliminate the buildup of condensation on the laser window are explored in an effort to improve the consistency of part quality, as well as to reduce maintenance requirements and costs.1.

BackgroundMechanical properties of laser sintered (LS) parts are primarily determined by the amount of thermal energy they absorb during each layer of the laser sintering process. Poor thermal distribution across the part bed is perhaps the largest contributor to inconsistency in part quality in LS, and has been the focus of study in recent research [1,2]. The research presented in this paper focuses on the second largest contributor to inconsistent part quality: the decrease of laser power and widening of the beam as gasses condense on the laser window. The resultant residue creates a significant barrier between the laser and the LS processing chamber, hindering the laser beam from passing through. A schematic of a laser sintering machine can be seen in Figure 1, where the laser is shown passing through the laser window in route to tracing the cross-section of a part at the part-bed.Figure 1.

A schematic of a laser sintering machine can be seen in Figure 1, where the laser is shown passing through the laser window in route to tracing the cross-section of a part at the part-bed.Figure 1. LS machine schematic; laser window and beamIn addition to the inconsistent part quality, the maintenance costs associated with cleaning and replacing laser windows are high. Traditionally, cleaning has been performed using alcohol and a soft cloth after each build. These repeated cleanings gradually scratch the surface of the window, eventually leading to its replacement. One laser window costs roughly $2,000, and each window has a lifespan of ~2-3 months when run on aerospace production machines at Harvest.The gasses that condense on the laser window are released, or outgassed, from the plastic powder as heat is applied. In the cases of PA 11 and PA 12, the two most common laser sintering materials, at least one of these gasses is known to be water vapor.

In the cases of PA 11 and PA 12, the two most common laser sintering materials, at least one of these gasses is known to be water vapor. As seen in Figure 2, the post-condensation reaction between two PA 12 monomers involves the release of a water molecule. This reaction is activated by thermal energy and aided by the flow of Nitrogen through the LS processing chamber, which keeps the chamber dry. The chemical makeup of other gasses released by various LS powders is presently unknown to the authors, although it is suspected that volatile monomers, oligomers, and powder additives (such as flow agents and whiteners) may be outgassed [3].Figure 2. Post-condensation reaction for PA 123D Systems currently utilizes the flow of heated Nitrogen below the laser window to hinder the ability of gasses to condense on the window. A strip heater is used to heat the Nitrogen piping to 90 °C prior to entering the chamber.

A strip heater is used to heat the Nitrogen piping to 90 °C prior to entering the chamber. However, at a flowrate of 7.5 LPM, the Nitrogen itself heats up to a more mild temperature of roughly 50 °C. The incoming Nitrogen flows through three outlets directed diagonally upward at the lower surface of the window. The authors suspect that this flow of Nitrogen has limited effectiveness because it is relatively cool and may even force outgassed material into the window.2. Effect of Residue on Laser PropertiesIn order to quantify the effects of window degradation on the properties of the laser at the part-bed surface, three test builds were run consecutively without cleaning of the window between builds. The builds were run at a laser power of 46 W using FR-106, a flame retardant, PA 11-based material from Advanced Laser Materials (ALM). The first two builds were geometrically identical, with a build height of 16” and a build time of 47 hours.

The first two builds were geometrically identical, with a build height of 16” and a build time of 47 hours. The third build was 6” high and ran for 16 hours. Before and after each build, laser power was measured in 5% increments of the maximum available laser power, 70 W. Additionally, beam intensity profiles were measured for three cases: no laser window, a clean laser window, and the laser window at the end of the third build. The plots in Figure 3 show the nominal and normalized laser power curves measured at the end of each of the three builds, as well as the nominal and normalized power curves before the start of the first build. Residue buildup appears to cause a percentage drop in laser power, rather than a constant drop. This is readily seen in the normalized plot, where, for example, the post-build 1 laser power is roughly 10% lower than the pre-build 1 laser power for nearly all of the laser power set-points.

This is readily seen in the normalized plot, where, for example, the post-build 1 laser power is roughly 10% lower than the pre-build 1 laser power for nearly all of the laser power set-points. For the third build, laser power did not drop significantly because the building time was low.Figure 3. Nominal and normalized power curves for the standard window housing designResidue buildup also affects the intensity profile of the beam. Table 1 summarizes the half-width values of the laser beam for the three previously mentioned cases. The values represent the averages of 250 measurements for each case. As shown Table 1, for lower laser powers such as 20 W, the intensity profile stays relatively unaffected by residue, but for higher laser powers such as 50 W, the intensity profile widens significantly with the presence of residue on the window.Table 1.

Half-Width Values for Various Beam Profiles20 W - No Window20 W - Clean Window20 W - Window After 3rd Build50 W - No Window50 W - Clean Window50 W - Window After 3rd buildHalf-Width in X-direction (μm)511.3471520.2568.4553734.9Half-Width in Y-direction (μm)506.1477.9510.2530.1550718.6Three-dimensional plots of the intensity profiles for the six cases can be seen in Figures 4 and 5. The three-dimensional plots are instantaneous images, and may not necessarily represent the average intensity profiles of the beam. For the higher wattage cases, the maximum power values appear to be clipped, which could cause errors in the half-width measurement. However, capturing the higher-power values would only serve to further decrease the half-width measurement for the clean and no-window cases, further separating them from the dirty-window case. The relative effect of residue buildup can still be visualized versus no window and a clean window for the 50 W case in Figure 5.Figure 4.

The relative effect of residue buildup can still be visualized versus no window and a clean window for the 50 W case in Figure 5.Figure 4. (a) 20 W, no window (b) 20 W, clean window (c) 20 W, window after 3rd buildFigure 5. (a) 50 W, no window (b) 50 W, clean window (c) 50 W, window after 3rd build3. Design to Reduce Buildup of CondensationDue to the thermal nature of the residue issue, a thermal approach has been taken to prevent the condensation of outgassed material on the laser window. The concept behind the design is straightforward: heat the laser window to a high enough temperature so as to prohibit any outgassed material from condensing on it. Although the background information regarding the types of gasses in question is not fully understood, experimental methods can be used to determine what temperature is necessary to prevent condensation.

Thus far, in the initial stages of research, a temperature of 125 °C has shown promising results.Since the upper surface of the laser window is exposed to ambient conditions, providing adequate heat to the window could prove difficult using the standard housing. Thus, a modified housing was designed such that two laser windows are utilized with a small pocket of air in between. Two 50 W heaters are placed in between the windows along with a thermocouple to provide feedback. A PID temperature controller is used to hold the temperature at a set value. CAD images of the design can be seen in Figure 6, below.Figure 6. CAD images of modified laser window housing4. Experimental ResultsThe first prototype was manufactured through laser sintering of glass-filled PA 12. This served as a form and fit prototype and provided an apparatus for testing thermal control.

This served as a form and fit prototype and provided an apparatus for testing thermal control. A 50 W heater was successful in heating air between the lenses to a temperature of at least 150 °C in an ambient environment; however, the prototype experienced melting, as seen in Figure 7, due to the high temperatures at the surface of the heater.Figure 7. First prototype of glass-filled PA 12 with melt-downA second prototype was machined out of Aluminum to be used in the LS machines. Due to the high thermal conductivity of Aluminum, a second 50 W heater was added to the secondprototype to provide sufficient energy input, and insulation was added to the interior surfaces. Images of the Aluminum prototype with heaters and thermocouple can be seen in Figure 8. An image of the Aluminum mounted in a LS machine and undergoing heating can be seen in Figure 9.Figure 8. Aluminum prototype with heaters and thermocoupleFigure 9.

An image of the Aluminum mounted in a LS machine and undergoing heating can be seen in Figure 9.Figure 8. Aluminum prototype with heaters and thermocoupleFigure 9. Aluminum prototype mounted in LS machine and undergoing heatingAs in the testing of the standard window housing discussed in Section 2, three consecutive builds were run using the modified housing without cleaning the laser window in between builds. The builds run were geometrically identical to the builds discussed in Section 2 as well: the first two builds were 16” tall and lasted 47 hours, while the third build was 6” tall and lasted 6 hours. The air between the two laser windows was heated to a set-point of 125 °C and held throughout the course of each build. The 125 °C set-point was chosen to ensure that any outgassed water vapor would be unable to condense on the window during the build.

The 125 °C set-point was chosen to ensure that any outgassed water vapor would be unable to condense on the window during the build. Additional testing at higher temperatures will be the focus of further research, because it is unknown whether other outgassed materials have condensation points above or below this temperature. Laser power curves were recorded before and after build and can be seen in Figure 10, below.Figure 10. Nominal and normalized power curves for the modified window housing designThe plots in Figure 10 show significant improvement in power loss compared to the standard window housing design. Whereas the standard design showed 10% power loss after the first build, the modified design shows anywhere from 0-3% power loss after the first build. Further comparison between the two designs is made in Figure 11, where the pre-build 1 and post-build 3 power curves for each design are overlaid.

Further comparison between the two designs is made in Figure 11, where the pre-build 1 and post-build 3 power curves for each design are overlaid. After three builds using the original design, a 20-25% power loss is recorded, whereas a 3-5% power loss is recorded for the modified design. A final visual comparison can be seen in Figure 12, where the window from the original design is noticeably dirty after the third build, while the window from the modified design is much cleaner after the third build.Figure 11. Standard design and modified design power curves overlaidFigure 12. Visual comparison of original and modified designs after third build5. Conclusions and Further WorkResidue buildup on the laser window during laser sintering causes significant changes in laser characteristics at the part-bed, decreasing power and widening the intensity profile of the beam.

Residue forms on the laser window when the feedstock materials outgas various substances which condense upon contact with the laser window. A new design was developed to heat the laser window so as to prevent condensation, and subsequently prototyped and tested with promising results. By holding the air between two laser windows to 125 °C and running a 47 hour build using ALM FR-106 material, much less power was lost than in the original design. After three builds and 112 hours of build time, the new design caused a decrease in power of 3-5%, whereas the original design caused a decrease in power of 20-25%.Further research is needed to identify the gasses present in the build chamber that condense on the window, so as to fully understand how to keep them from condensing. In addition, further testing will be performed at higher temperatures in an effort to prevent gasses with higher condensation points from forming residue on the window.References:[1] M. Yuan and D.

These components are often subjected to multiaxial stress states due to their typically complex design accompanied by residual stresses and/or multiaxial external loading. Therefore, understanding the fatigue behavior of AM materials under multiaxial-type loadings is necessary for ensuring reliable in-service component performance. In this study, the effects of layer orientation on the multiaxial fatigue behavior of Ti-6Al-4V fabricated via a laser-powder bed fusion (L-PBF) process was investigated. Tubular thin-walled multiaxial specimens were fabricated in vertical and diagonal orientations with respect to the build plate. Specimens were tested under axial, torsional, in-phase axial/torsional, and 90° out-of-phase axial-torsional cyclic loadings.

Specimens were tested under axial, torsional, in-phase axial/torsional, and 90° out-of-phase axial-torsional cyclic loadings. Upon failure, the crack orientation of vertical and diagonal specimens was correlated to the type of loading, which illustrated the failure mechanism of L-PBF Ti-6Al-4V and justified the variations in the fatigue lives of specimens.Keywords: Additive Manufacturing, Failure Mechanisms, Multiaxial Fatigue, Titanium.IntroductionAt its core principle, additive manufacturing (AM) technology allows for the incorporation of 3D computer aided designs into solid parts by melting and joining the feedstock material in a layer-by-layer manner [1–3]. The most commonly used AM process is laser-based powder bed fusion (L-PBF) [3–7], whose process consists of applying a layer of metallic powder over a substrate and melting the powder in a pre-determined section to create the first layer of part. This process is repeated until the part is completed [3].

This process is repeated until the part is completed [3]. Some inherent disadvantages of L-PBF process, also found in other metal AM techniques, include rough surface finish and the presence of tensile residual stresses, as well as internal defects such as lack of fusion (LOF) and porosity [6–8]. Microstructural characteristics, such as grain size, morphology, and orientation, are also influenced by the thermal history (e.g. cooling rate, reheating, etc.) experienced by parts during fabrication [7,8]. In addition, depending on the layer orientation (i.e. the build orientation with respect to a1384build plate), parts can experience different thermal histories, which subsequently affect morphology and direction of manufacturing-induced defects [7,9,10].Recent studies [9–12] have focused on the effects of layer orientation on mechanical properties of several AM materials under uniaxial quasi-static and fatigue loadings. Bača et al.

Bača et al. [10] investigated the fatigue behavior of direct metal laser sintering Ti-6Al-4V, and demonstrated that for all stress amplitude levels employed, the vertically-fabricated specimens had considerably shorter fatigue resistance than the horizontal specimens. However, several studies have found minimal effect of the layer orientation on mechanical behavior of AM metals [13,14]. For example, Torries et al. [13] investigated the effect of layer direction on the fatigue performance of L-PBF Ti-6Al-4V, and reported similar results for the horizontal, vertical and diagonal specimens under the same surface conditions (i.e. polished). Minimal differences in fatigue lives were attributed to low defect density. Based on these findings, it is suggested that further studies regarding layer orientation effects under more complex loadings are necessary.Due to complex part geometry required for given applications, AM metallic parts are often subjected to multiaxial cyclic loading.

In some cases, even under simple uniaxial loading, parts may experience multiaxial state of stresses. Fatemi et al. [6] recently investigated the effects of surface roughness (i.e. as-built and machined) on the multiaxial fatigue behavior of L-PBF Ti-6Al- 4V under axial, torsion, and in-phase (IP) loadings. It was shown that surface machining improved the overall fatigue resistance of L-PBF Ti-6Al-4V under the aforementioned loadings. It was also found that the fatigue data for different surface roughness conditions correlated well with the maximum principal stress failure criterion, due to the materials brittle behavior. However, failure on the maximum principal stress or strain plane is not typical in titanium alloys. Previous studies [6,15–17] have demonstrated that failure of wrought titanium alloys is dominated by the maximum shear stress or strain plane. This failure mode is often explained by their ductile behavior allowing cracks to nucleate from slip planes.

This failure mode is often explained by their ductile behavior allowing cracks to nucleate from slip planes. For brittle materials, such as L-PBF Ti-6Al-4V, cracks tend to originate from un-melted powder particles, LOF, or porosity [6,18,19].To the best of the authors’ knowledge, no research has been conducted on the multiaxial fatigue behavior of AM parts considering the influence of layer orientation. Therefore, in this study, the effects of layer orientation on the microstructural characteristics, mechanical failure modes, and multiaxial fatigue behavior of Ti-6Al-4V fabricated via L-PBF are investigated. The loadings include under axial, torsional, IP, and 90° out-of-phase (OP). Major findings and fatigue life correlation based on different failure criteria are presented.Experimental ProcedureGas-atomized Ti-6Al-4V Grade 23 powder with a particle size of 15-45 µm, manufactured by LPW Technology, was used to fabricate specimens via L-PBF process using an EOS M290 system.

The default process parameters were employed, which include 280 W laser power, 1200 mm/s travel speed, 100 mm stripe width, 140 µm hatch spacing, 67° hatch rotation, and 30 µm layer thickness. The size and geometry of the tubular multiaxial thin-walled specimens, designed in accordance to ASTM E2207-15 standard [20], is depicted in Fig. 1. Specimens were fabricated in vertical, and diagonal (i.e. 45° from to the build plate) orientations. Stress relieve heat treatment was performed at 704 °C (1300 °F) under argon atmosphere for 1 hour followed by furnace cooling. Specimens were then removed from the build plate, and their gage section was lightly polished to remove residual powder prior to testing. No post-processing machining was carried1385out on these specimens. The microstructure was characterized by cutting the gage section of specimens in the longitudinal and radial directions.

No post-processing machining was carried1385out on these specimens. The microstructure was characterized by cutting the gage section of specimens in the longitudinal and radial directions. In the case of the diagonal specimens, the longitudinal dissection was performed along the supported side (i.e. down-skin). The cut samples were then polished and etched with Kroll’s reagent.Fig. 1. Tubular multiaxial specimen with uniform gage section per ASTM standard E2207-15 [20]. All dimensions are in mm.Monotonic and fatigue tests were conducted using a MTS 809 axial/torsion closed-loop servohydraulic test system equipped with 100 kN axial force, and a 1100 N m torsional force load capacity. Displacement- and rotation-controlled mode at 0.1 mm/min and 0.1 °/min were used in monotonic tension and torsion tests respectively. An axial/torsion mechanical extensometer was used to record strain values.

An axial/torsion mechanical extensometer was used to record strain values. Fully-reversed (R = -1) fatigue tests were carried in force and torque control modes using a tapered sinusoidal waveform signal. Force and torque control modes were used due to the fatigue deformation response of the material being fully elastic. The loading paths employed in this study were axial, torsional, IP, and OP. Depending on the type of loading, test frequency was adjusted (0.4 - 2 Hz) to minimize any strain rate effects. A duplicate test for each loading configuration was conducted to verify the consistency of the collected data.Results and DiscussionMonotonic Deformation The monotonic axial and torsional stress-strain responses of the vertical and diagonal specimens are presented in Figs. 2(a) and 2(b), respectively. Strain was recorded up to the extensometer limits (i.e. curves ended with arrows); the extensometer was then removed and the test continued until final fracture.

2(a) and 2(b), respectively. Strain was recorded up to the extensometer limits (i.e. curves ended with arrows); the extensometer was then removed and the test continued until final fracture. As shown in these figures, the layer orientation, with respect to the loading, influences the mechanical response of L-PBF Ti-6Al-4V.Fig. 2(a) shows that the diagonal specimens have slightly higher yield stress (~ 55 MPa) as compared to vertical specimens, while both specimens seem to exhibit similar work hardening rates, retaining their near constant gap of 60 MPa. Furthermore, vertical specimens appeared to be more ductile as the displacement at final fracture was ~17% higher than the diagonal specimens (the values are not presented in the figures). Similar findings were reported for selectively laser melted and direct laser deposition AM Ti-6Al-4V studies [8,9,21], where it was concluded that1386layer orientation being parallel to the load (e.g.

2. Monotonic stress-strain curves of L-PBF Ti-6Al-4V specimens under (a) axial and (b) torsion loading.As is seen in Fig. 2(b), the torsional behavior is also affected by the layer orientation, where the vertical specimen had slightly higher yield and ultimate shear strengths than the diagonal ones by ~65 MPa. The vertical specimens had increased shear ductility due to final fracture, having1387~10% greater angle of rotation than the diagonal specimens. In Fig. 2(b), the torsional behavior of the vertical and diagonal specimens are also compared to a recent study for L-PBF Ti-6Al-4V [19]. It can be seen in this figure that the mid-section shear stress-strain responses in [19] are similar to the diagonal specimen in the present study, but with significantly less ductility.

It can be seen in this figure that the mid-section shear stress-strain responses in [19] are similar to the diagonal specimen in the present study, but with significantly less ductility. This could be attributed to the different AM system and process parameters used for specimen fabrication.The anisotropy in elongation to failure under both axial and torsional loadings was consistent with the existing literature, where loading direction being parallel to the columnar prior β grains (e.g. vertical specimen under monotonic axial loading) resulted in higher elongation [21]. Epitaxial grain growth orientation of the vertical and diagonal specimens is presented in Figs. 3(a) and 3(b), respectively. The observed anisotropy in ductility may be due to the formation of primary α on prior β grain boundaries, which are subjected to accelerated damage under larger principal tensile stress (i.e. mode I or crack opening mode) component for diagonal specimens.

mode I or crack opening mode) component for diagonal specimens. The variations in yield stress and ultimate tensile strength (i.e. ~5% variance) may be justified by lath thickness differences and overall texture of the specimens stemming from their thermal , 𝛼𝛼 histories [23]. The variations in thermal histories, i.e. cooling rates, is due to the changes in the thermal conduction capacity or volume of the material underneath the molten metal layer caused by sample orientation with respect to the substrate [7].n o i t c e r in o i t c e r iD d l i u BD d l i u B(a)(b)Fig. 3. Microstructure of L-PBF Ti-6Al-4V built in (a) vertical and (b) diagonal orientations.Fatigue Behavior Two equivalent stress amplitude levels (331 and 234 MPa) based on von Mises equivalent stress were selected for fatigue testing as the results were expected to be in between the short and long cycle regime.

The stress-life results of vertical and diagonal specimens under axial, torsional, IP, and OP loading tests are presented Figs. 4(a) and 4(b). It is important to emphasize that the shear stress at the surface was used to analyze the fatigue behavior of torsion and axial/torsional tests. As seen in Fig. 4(a), when using von Mises equivalent stress criterion for axial, torsional, IP, and OP loadings, the fatigue lives were expected to be similar given that an equivalent stress amplitude of 331 MPa was used. However, results show a fatigue life variance with a factor of 26 when all loadings are considered. This is attributed to this criterion typically providing better1388fatigue life correlation for ductile materials, and L-PBF Ti-6Al-4V having a brittle behavior and tensile failure mode [6,18]. This also agrees with the findings by Fatemi et al. [6], and the final crack orientation of the specimens being oriented along the maximum principal plane.

This also agrees with the findings by Fatemi et al. [6], and the final crack orientation of the specimens being oriented along the maximum principal plane. Further explanation on the damaging effects of OP loading can be found in [18].1,000Vertical Diagonal) a P MIP OP Axial Torsion TorsionIP OP Axial(a σ, s s e r t St n e l a v i u q E1002.E+042.E+05 Reversals to failure, 2Nf2.E+06(a)1,000Vertical Diagonal) a P M(IP OP Axial Torsion TorsionIP OP Axial1 a σ ̄ , l a p i c n i r P m u m i x a MFit + 3 - 3 R2 = 0.721002.E+042.E+05 Reversals to failure, 2Nf2.E+06(b)Fig. 4. Correlation of fatigue test results of L-PBF Ti-6Al-4V using (a) von Mises equivalent stress, and (b) maximum principal stress criterion.Under OP loading and an equivalent stress amplitude of 331 MPa, it can be seen that the diagonal specimens had longer fatigue life than the vertical specimens; while, for torsional loading, the results were vice versa.

For OP loading and an equivalent stress of 234 MPa, the effects of layer orientation were negligible. For IP and axial loadings, fatigue lives were similar for the1389diagonal and vertical specimens. Differences between OP and torsional, IP and axial, at 331 MPa stress level can be attributed primarily to the layer orientation with respect to the maximum principal stress plane under each loading condition. For OP loading, the maximum tensile and maximum shear stress plane is oriented at 0°. Therefore, in the case of vertical specimens, the layers are oriented along the maximum tensile stress plane, thus the stress concentration due to surface roughness causes the crack to propagate directly across the grains without much resistance. This in turn, may shorten the fatigue life remarkably.

This in turn, may shorten the fatigue life remarkably. In the case of torsional loading, the layers of the diagonal specimens are oriented along the maximum tensile stress plane; therefore, the stress concentration due to the surface roughness will cause failure to occur along this plane. The maximum tensile stress plane is also oriented at 0° or 90° with respect to the prior β grain boundaries, which may also contribute to the reduced fatigue life. This is due to the primary α on the prior β grain boundaries exhibiting less resistance to crack propagation when the maximum tensile stress is perpendicular to the prior β grain boundaries [21].Since it was determined that failure occurred along the maximum principal plane (i.e. tensile mode), the maximum principal stress amplitude was correlated to fatigue life, as presented in Fig. 4(b).

tensile mode), the maximum principal stress amplitude was correlated to fatigue life, as presented in Fig. 4(b). As can be seen in this figure, the principal stress versus fatigue life is appropriately correlated and the data falls within the scatter bands of a factor of three from the fit. Although the layer orientation is influential in terms of fatigue performance, the maximum principal stress criterion can address the observed discrepancies. It can also be observed that, under OP loading at 234 MPa stress amplitude level, the data falls on the lower scatter band, which indicates that the maximum principal stress may not be the appropriate criterion for L-PBF Ti-6Al-4V.

Therefore, further studies are needed to obtain more appropriate critical plane criterion to improve the overall results.ConclusionsThis study examined the effects of layer orientation on the monotonic deformation and fatigue behavior of L-PBF Ti-6Al-4V under several types of loading including axial, torsional, in- phase (IP) and 90° out-of-phase loading (OP) axial-torsional. The following conclusions can be inferred from the experimental results and analysis:1. Comparison of vertical and diagonal specimens under monotonic axial and torsional loadings indicated the deformation and failure behavior of L-PBF Ti-6Al-4V is affected by the layer orientation (i.e. microstructure and defect orientation) with respect to the maximum principal plane. Implying that, when the maximum principal loading is parallel to the grains, the stress response will decrease and ductility will increase.2.

Implying that, when the maximum principal loading is parallel to the grains, the stress response will decrease and ductility will increase.2. Under different combination of axial/torsional fatigue loadings, the layer orientation plays an important role in the resultant fatigue resistance. For example, diagonal specimens offered superior results under OP loading when compared to the vertical specimens, while vertical specimens outperformed the diagonal ones under torsional loading.3. The von Mises stress criterion was not able to correlate fatigue results as the failure mode of L-PBF Ti-6Al-4V was found to be along the maximum principal plane, and not the maximum shear stress plane.4.

The maximum principal stress criterion presented acceptable fatigue life correlation due to the brittle behavior of L-PBF Ti-6Al-4V, where all the data was correlated within a scatter factor of three.1390AcknowledgementsThis research was partially funded by the National Science Foundation (NSF) under grantNo.1657195References [1] Weber C, Peña V, Micali M, Yglesias E, Rood S, Scott JA, et al. The role of the national science foundation in the origin and evolution of additive manufacturing in the United States. Sci Technol Policy Inst 2013;1.[2] Frazier WE. Metal additive manufacturing: A review. J Mater Eng Perform 2014;23:1917–28.[3] Thompson SM, Bian L, Shamsaei N, Yadollahi A. An overview of direct laser deposition for additive manufacturing; Part I: Transport phenomena, modeling and diagnostics. Addit Manuf 2015;8:36–62.[4] Lewandowski J, Seifi M.

Int J Fatigue 2010;32:1862–74.[16] Wu Z-R, Hu X-T, Song Y-D. Multiaxial fatigue life prediction for titanium alloy TC4 underproportional and nonproportional loading. Int J Fatigue 2014;59:170–5.[17] Nakamura H, Takanashi M, Itoh T, Wu M, Shimizu Y. Fatigue crack initiation and growth behavior of Ti–6Al–4V under non-proportional multiaxial loading. Int J Fatigue 2011;33:842–8.[18] Fatemi A, Shamsaei N. Multiaxial fatigue: An overview and some approximation models forlife estimation. Int J Fatigue 2011;33:948–58.[19] Fatemi A, Molaei R, Sharifimehr S, Shamsaei N, Phan N. Torsional fatigue behavior of wrought and additive manufactured Ti-6Al-4V by powder bed fusion including surface finish effect. Int J Fatigue 2017;99:187–201.[20] ASTM E2207-15 Standard practice for strain-controlled axial-torsional fatigue testing with thin-walled tubular specimens, ASTM International, West Conshohocken, PA, 2015. [21] Carroll BE, Palmer TA, Beese AM.

[21] Carroll BE, Palmer TA, Beese AM. Anisotropic tensile behavior of Ti–6Al–4V components fabricated with directed energy deposition additive manufacturing. Acta Mater 2015;87:309– 20.[22] Carrion PE, Shamsaei N, Daniewicz SR, Moser RD. Fatigue behavior of Ti-6Al-4V ELIincluding mean stress effects. Int J Fatigue 2017;99:87–100.[23] Sridharan N, Chaudhary A, Nandwana P, Babu SS. Texture evolution during laser directmetal deposition of Ti-6Al-4V. JOM 2016;68:772–7.1392.

fects of local fiber orientation state on thermal-mechanical behaviors of composite parts made by large area polymer deposition additive manufacturing Zhaogui Wang1, Zhenyu Fang1, Douglas E. Smith2 1Department of Mechanical Engineering, Dalian Maritime University, Dalian, China, 116000 2Department of Mechanical Engineering, Baylor University, Waco, TX, USA, 76798 AbstractShort carbon fiber enhances the dimensional stability and material strength of composite parts created via large area polymer deposition additive manufacturing, which has been used for rapid fabrications of large-dimension composite parts and tooling. Nevertheless, the flow-induced fiber orientation formed during the material extrusion and deposition leads the deposited composites exhibit non-homogeneous thermal-mechanical behaviors. This study evaluates the fiber orientation state of a 20 wt.% CF-PEI composite fabricated by polymer deposition using the fully coupled flow/orientation approach.

This study evaluates the fiber orientation state of a 20 wt.% CF-PEI composite fabricated by polymer deposition using the fully coupled flow/orientation approach. The material properties are computed by considering the deposited bead as heterogeneous segments with different local fiber orientation states. The heterogeneous thermal conductivity and expansion coefficient exhibit maximum local differences of 29% and 21%, respectively. The orientation-homogenized material properties are implemented to the finite element simulation for a large area additive manufacturing process of a single bead and notable differences are seen between results computed by employing the homogenous and heterogeneous properties. IntroductionLarge Area polymer deposition Additive Manufacturing (LAAM) is seen rapid growth in the area of composite manufacturing due to the high efficiency enabled by the screw-extruder-based material feeding system [1].

Large-dimension composites parts and tooling including the full-size car, boat or wind turbine blade mold are 3D printed directly through the LAAM technology with less cost in time and materials as compared to conventional methods [2-5]. The carbon fiber filled composites provide superior stiffness and low thermal expansion that enhance the mechanical performance of fabricated structures. Specifically, the discontinuous fibers re-orient as the material flow being extruded and deposited onto the material substrate and the resulting fiber orientation in the beads can be quite different locally as seen in Figure 1. Nixon, et al. evaluated the internal geometric effects of the extrusion nozzle on the fiber orientation alignment using Moldflow software (Moldflow Corporation, Framingham, MA, USA), where a convergent nozzle geometry was found to yield higher fiber alignment than a divergent designed nozzle [6]. Heller, et al.

Heller, et al. employed an axisymmetric 2D flow model via COMSOL (COMSOL Inc., Burlington, MA, USA) simulating the nozzle flow with a short section of free extrudate included, and found that the die swell at nozzle exit reduced the principal fiber orientation by ~20% [7]. Extended works based on Heller’s model explored the non-Newtonian fluid rheology effects [8] and the screw swirling motion [9] on the resulting fiber orientation of an extrudate composite. Above mentioned studies are done using a weakly coupled formulation between the flow kinematics and fiber orientation evolution, where the presence of fibers is ignored during theSolid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference1389computation of the flow kinematics. In contrast, Beversta, et al.

In contrast, Beversta, et al. employed smoothed particle hydrodynamics method simulating the transient flow development during the planar deposition process [10], where the flow kinematics and fiber orientation were solved in a mutually coupled formulation. Under the same fully coupled scheme, Wang and Smith, on the other hand, applied the finite element approach evaluating the flow and fiber orientation fields in a nozzle-extrudate flow [11] and planar deposition flow [12], where a quasi-steady state of fiber orientation in the extruded composites was obtained for further material properties estimations. Computed results from prior studies (e.g., [7,8,11,12]) indicated that the fiber orientation pattern within the deposited beads has a direct effect on the material properties of the printed parts. Figure 1. Complex local orientation across a composite bead extruded by LAAM.

Figure 1. Complex local orientation across a composite bead extruded by LAAM. Particularly, this study is interested in the effect of local fiber orientation within a bead on the material properties of the printed composites through LAAM systems. Firstly, the non-Newtonian flow kinematics and fiber orientation fields in a planar deposition flow domain are solved in a fully coupled formulation via the finite element method. The flow field is modeled using the power law model and the fiber orientation state is evaluated using the Advani-Tucker orientation tensor approach through the Folgar-Tucker fiber orientation evaluation equation with isotropic rotary diffusion and orthotropic fitted closure. Then, the material properties of a deposited bead are estimated through the orientation homogenization approach, including the elastic constants, coefficients of thermal expansion and thermal conductivity.

Herein, we imposed a fully developed fiber orientation state reported by Heller, et al. [7], assuming that the orientation state reaches somewhat a steady state at the inlet of the flow. The fiber aspect ratio is set as 15, in an averaged sense, which is a safe estimation based on related experiments (see, e.g., [18,19]). The  is set as 0.0055, based on Equation 8.1392Figure 2. Boundaries of the flow domain of interest. The flow velocity fields vx, vy are plotted in Figures 3 and 4, respectively. It is seen that the velocity profile vx reaches an equilibrium roughly in the middle of the deposited bead. The fiber orientation tensor fields of A11 and A22 components are plotted in Figures 5 and 6, respectively. It is clearly seen that the fibers are highly aligned along the direction of deposition. The fiber orientation state also reaches a steady state much faster as compared to the flow kinematics, comparing the contours of Figures 3 and 5.

The fiber orientation state also reaches a steady state much faster as compared to the flow kinematics, comparing the contours of Figures 3 and 5. Furthermore, we employed the steady state fiber orientation seen in the flow end to evaluate the material properties through the orientation homogenization method [16]. The second order fiber orientation tensor diagonal components at the flow end across the bead thickness are given in Figure 7. The constituent materials properties of the carbon fiber and PEI matrix are given in Table 1. The computed material properties at each location of the computed fiber orientation tensor (cf. Figure 7) are ultimately integrated across the bead thickness direction to obtain the effective properties of the bead. Computed properties including elastic constants, thermal conductivity and expansion coefficients are given in Table 2. Note, due to the high principal fiber alignment (cf.

Fiber orientation tensor components at the flow end across the bead thickness. Table 1. Material properties of the phase materials of a 20 wt.% CF-PEI (~16 vol.%). Material E (GPa)   22 5.19 N/A N/A 8.302e-5 0.144 33 1.86 N/A N/A 7.079e-5 0.182 12 N/A 1.86 0.19 N/A N/A 23 N/A 1.62 0.52 N/A N/A 13 N/A 2.19 0.21 N/A N/ACarbon fiber 230 0.2 -2.6e-6 3.06It is noticed that the steady state fiber orientation at the flow end exhibit notably difference locally, where, e.g., the maximum A11 component (i.e., 0.9568) is 59% higher than that of the minimum (i.e., 0.6011). Hoskins, et al. measured the coefficients of thermal expansion on LAAM-made ABS composites and results show a similar trend as seen in A11 components appearing in Figure 7. In simulating the residual stress of LAAM-made parts, they partitioned the deposited strand in five subdomains with five different material properties estimated based on their measured data [20].

In simulating the residual stress of LAAM-made parts, they partitioned the deposited strand in five subdomains with five different material properties estimated based on their measured data [20]. Herein, a similar strategy is employed to define our deposited PEI composites for further printing process simulation, where the bead is divided into five equal-length sections along the direction of bead thickness. Material properties of each sections are estimated using the local fiber orientation within that section via the same orientation homogenization approach applied above [16]. The entire material properties estimated considering the local fiber orientation state are given in Table 3 for further simulations. Specifically, we plot three properties, for instance, through Figure 8, where the properties of E11, 11, .are included. It can be seen that the locally defined material properties exhibit dramatically differences as compared to the uniformly homogenizedTable 2.

The unit extrusion element is in 3mm thickness as normally seen for LAAM-deposited beads [11,12]. The width of the unit is assumed as 8mm. The print speed (i.e., nozzle moving speed) is set as 100 mm/s (see, e.g., [11]), and the step analysis time for a unit element is 0.005 second. Thermal and mechanical boundary conditions are set based on a related preceding study [21]. Specifically, the initial temperature of each extrusion unit is 380 ℃, which is common for CF-PEI [18]. The temperatures of the substrate and ambient air are 90 ℃ and 110 ℃, respectively. The air convection coefficient is set as 90 /( ). Computed Mises stress contours by employing uniformly homogenized material properties and five-layer locally homogenized material properties are given in Figures 9 and 10, respectively.Finite Element Simulation of LAAMLayer 1 11 13.55 N/A N/A 2.878e-5 0.551Table 3. Locally homogenized material properties of deposited CF-PEI composites.

Locally homogenized material properties of deposited CF-PEI composites. Layer Direction E (GPa) G (GPa)   1397The stress contours in Figures 9 and 10 are in similar fashion, while the maximum stress values of the simulations exhibit a notable difference. By employing locally homogenized material properties, the maximum stress reduces 14% as compared to that computed using uniformly homogenized material properties. In addition, we present the Mises stress in the direction of material loading in Figures 11 and 12, where significant differences are seen clearly in the contours, especially across the bead thickness. This is a result of the high degree of material anisotropy as seen in Tables 2 and 3. Further, the maximum deformation of the as-printed bead in uniform properties and localized properties simulations are 5.64E-3 mm and 5.13E-3 mm, respectively, i.e., ~10% difference for the LAAM process of a 100mm-length single bead.

From where it is expected that for large-dimension LAAM parts, the over-estimation in material deformations may be much higher by using uniformly homogenized properties over locally homogenized properties. To this end, the local orientation effects are considered as a crucial component in simulations of LAAM composites manufacturing. Figure 9. Mises stress contour of a 20% CF-PEI to form a 100mm-length single strand by employing uniformly homogenized material properties. Figure 10. Mises stress contour of a 20% CF-PEI to form a 100mm-length single strand by employing five-layer homogenized material properties.1398Figure 11. Mises stress component S11 contour of a 20% CF-PEI to form a 100mm-length single strand by employing uniformly homogenized material properties. Figure 12. Mises stress component S11 contour of a 20% CF-PEI to form a 100mm-length single strand by employing uniformly homogenized material properties.

Figure 12. Mises stress component S11 contour of a 20% CF-PEI to form a 100mm-length single strand by employing uniformly homogenized material properties. ConclusionA two-step finite element based method is employed to simulation the LAAM process of 20 wt.% carbon fiber filled PEI polymer. The extrusion deposition process is simulated via a fully coupled flow/orientation approach, where the flow kinematics and fiber orientation tensor fields are solved. Employing the steady state fiber orientation tensor at the deposition flow end into the orientation homogenization method, the elastic constants, thermal conductivity and thermal expansion coefficients of the deposited CF-PEI composites are estimated. The fiber orientation effects on the material properties are considered in the property estimations, where the deposited bead is equally divided into five layers and fiber orientation within each layer are homogenized independently.

The thermal-mechanical transient history of the LAAM process of a 100mm-length1399single deposited strand is simulated via ABAQUS, employing the computed material properties. As compared to results computed by employing uniformly homogenized material properties, the locally homogenized material properties simulation exhibits significant difference in the Mises stress contours, where the maximum stress value reduces 14%. In addition, the uniformly homogenized properties simulation over-estimates the maximum deformation in the deposited bead as compared to the localized properties simulation. To these ends, the fiber orientation effects on the material properties are considered as an important component in LAAM simulations. AcknowledgementsThe authors would like to appreciate the support from Dalian Maritime University and Baylor University. This work was supported by the Fundamental Research Funds for the Central Universities of China (No. 3132021337).

fects of Oxygen Inhibition and Post-Processing on Exposure Controlled Projection Lithography Process AccuracyYing Zhang, Amit Jariwala*, David W. RosenGeorge W. Woodruff School of Mechanical EngineeringGeorgia Institute of TechnologyAtlanta, Georgia, 30332Corresponding author. Tel.: +1 404 894 3931 Email: Amit.jariwala@gatech.eduAbstracta mask-projection stereolithography process which can be used to create micro lenses on flat or curved substrates. In the ECPL process, the ultraviolet light patterned by the dynamic mask passes through a transparent substrate to cure the photopolymer resin to a certain shape. The dimensions of the part can be controlled by the exposure time and functional pixels in the dynamic mask. In this paper, a modified process planning method is presented with the considerations of post- processing and oxygen inhibition, which can vary part dimensions significantly. The effects of post-processing and oxygen inhibition are studied and characterized.

The effects of post-processing and oxygen inhibition are studied and characterized. The accuracy of the lateral and vertical dimensions of the cured part are improved by the revised method. Experimental validation is obtained by fabricating samples using the ECPL system.Exposure Controlled Projection Lithography(ECPL)is1. IntroductionExposure Controlled Projection Lithography (ECPL) is an additive manufacturing process where components are solidified in a photopolymer resin by ultraviolet light. The ECPL process is a similar system as the Masked Projection Stereolithography process (MPSLA) [1] except that the curing of the photopolymer resin occurs by passing light energy through a transparent substrate. With control of the light intensity profile and curing time, the vertical in shape of the sample will be fabricated Stereolithography (SLA) process. The dimension of the sample can be controlled by the exposure time and functional pixels in the projector.

The dimension of the sample can be controlled by the exposure time and functional pixels in the projector. To project functional pixels, a series of binary bitmap images are used. Recently researchers such as Erdmann et al. [2] and Mizukami et al. [3] have developed similar techniques as the ECPL process. However, the knowledge of controlling the process to achieve high accuracy and precision in the final cured shape was not presented.instead oflayer-by-layer approach usedThe revised process planning method to control the lateral dimensions of the cured part in the ECPL process with consideration of a ray tracing model and photopolymerization kinetic model was presented by Jariwala et.al.[4]. The cured results showed a significant deviation with simulation results in dimensions. Therefore, a hypothesis was developed by Jariwala et.al [5] that the effect of oxygen diffusion and inhibition should be considered during the polymerization process.

Therefore, a hypothesis was developed by Jariwala et.al [5] that the effect of oxygen diffusion and inhibition should be considered during the polymerization process. The revised process planning method provided better control of the curing results.However, the accuracy was still not satisfied and the model was not suitable for all cases. In order to improve the accuracy, a similar process planning method with revised rate constant was presented by Jariwala et.al [5]. A unique process planning approach was demonstrated which simulated the empirical models in the COMSOL software package and then formulated in MATLAB software to simulate the entire curing process. With the experimentally validated rate constant, the error between simulation and experiments was still around 15%.In this paper, rate constants were revised again with experimental validation. The irradiance model was modified to fit the experimental data.

The irradiance model was modified to fit the experimental data. Furthermore, the effect of post- washing process was studied and reduced to an insignificant level. The revised process planning approach is presented and results demonstrate significant improvements in part dimensional accuracy.2. ECPL Systemthe block diagram of the ECPL process. As the block diagram indicates, the process model of the ECPL system contains three main parts: apply energy, process and post-process. To apply energy, the ultra violet light source apply through the beam expander to expand the laser beam diameter. Then the light passes through the diffuser to enlarge the diameter and homogenize the beam’s intensity and the collimating lens to collimate the light. After that, the UV coated mirror will be used to direct the light on a dynamic mask, which is the DMD projection system. The projection system is used to enlarge or reduce the image presented on the DMD and project it on the resin chamber.

The projection system is used to enlarge or reduce the image presented on the DMD and project it on the resin chamber. The part is cured in the resin chamber, which consists of two glass slides and a spacer which is shown in Figure 2. The uncured monomer resin is loaded inside.Figure 1 showsFigure 1：Block diagram of ECPL system [9]3. Existing ECPL Process Background3.1. Irradiance ModelThe ECPL system is analytically modeled Irradiance Model and irradiancein Photopolymerization Model. The model models the irradiance received by the resin in terms of the process parameters and is presented in detail in Jariwala et al. [6]. The UV Light energy is the primary energy source in the ECPL system. The total amount of energy is controlled by the exposure time and by the size of the irradiated region, given by the size of the projected bitmaps. The exposure (units of energy/area) is referred as the total amount of UV energy incident per unit area.

The exposure (units of energy/area) is referred as the total amount of UV energy incident per unit area. The total dose of energy provided to the resin is regulated by the distribution of irradiance at the substrate level. The relationship between the irradiance distribution on the DMD chip and the substrate of the resin chamber was investigated as the irradiance model in Jariwala et al. [6]. The irradiancetwo phases –Figure 2: Schematic of the resin chamber [9]model can be validated by placing a camera at the substrate level. Figure of the 3 shows simulation the results irradiance model with experimental data (irradiance detected by the camera). The result demonstrates irradiance model can that estimate the actual exposure profile on the substrate level though there were small deviations, which might be due to the misalignment of optics or the resolution difference between the simulation model and UV CCD camera.comparisonfromthe3.2.

Polymerization ModelFigure 3: Comparison of simulation results from Irradiance model (solid lines) with data from Camera (dashed lines). Red corresponds to 10pixels; green to 60pixels and blue to 90pixels [9]Photopolymerization is thereaction of monomers or macromers to produce solid polymeric structures by light-induced initiation and polymerization [7]. Most Stereolithography (SL) resins contain acrylate monomers. For an acrylate resin system, the usual catalyst is a free radical. In Stereolithography, the radical is generated photo chemically. The source of the photo chemically generated radical is a photo initiator, which reacts with an actinic photon. This produces radicals that catalyze the polymerization process.

The source of the photo chemically generated radical is a photo initiator, which reacts with an actinic photon. This produces radicals that catalyze the polymerization process. According to Beer-Lambert’s law of absorption, the exposure (mJ/cm2) decreases exponentially with depth [8].E(z) = 𝐸𝑚𝑎𝑥𝑒−𝑧 𝐷𝑝(1)where Dp is the resin “penetration depth” (a resin parameter) at the given wavelength and Emax is the exposure at the surface of the resin (z = 0). Z is the cured part height. Based on experimental observations, this model was modified in [3, 4] as follows:z ≈ 𝐷𝑝𝑠 ∗ ln⁡(𝐷𝑝𝐿 𝐷𝑃𝑆∗𝐸 𝐸𝑐+ 1 −𝐷𝑝𝐿 𝐷𝑃𝑆)(2)where Ec is the critical exposure (exposure threshold that induces photopolymerization), DPL is the penetration depth through liquid resin, and DPS is the penetration depth through solid resin. The parameters Ec, DpL and DpS are usually fit to experimental data at a specific resin composition and cure intensity.

The parameters Ec, DpL and DpS are usually fit to experimental data at a specific resin composition and cure intensity. Moreover, this empirical model was found not to be adequate to predict the actual reaction inside the resin. The diffusion effects of oxygen were investigated to have significant effects on the size, shape and properties of parts fabricated by Stereolithography [9] and were incorporated into the kinetic model of the ECPL system. The model incorporates the chemical reaction inside the resin with oxygen diffusivity. The polymerization kinetic model is presented briefly as follows. The concentrations of photoinitiator [In], radicals [R∙], unreacted double bonds [DB], and oxygen [O2] were modeled in the kinetic model. The reactionsconsidered by them were as follows [10].

The concentrations of photoinitiator [In], radicals [R∙], unreacted double bonds [DB], and oxygen [O2] were modeled in the kinetic model. The reactionsconsidered by them were as follows [10]. When the photopolymer resin receives light energy, the photoinitiator absorbs it and decomposes into two radicals with first order rate constant of Kd,lnKd → 2R ∙(3)The initiator decomposition rate, Kd, is well known in literature and is modeled as a function of the local intensity, which varies with depth (following the Beer-Lambert Law) [11]𝐾𝑑 =2.3Ф𝜀𝜆 𝑁𝐴ℎ𝑐𝐼0𝑒(−2.3𝜀[𝐼𝑛]𝑧)(4)where 0<ϕ<1 is the quantum efficiency of the photoinitiator, NA is Avagadro’s number, h is Planck’s constant, and c is the speed of light. The molar absorptivity of the resin, ε, depends upon the source wavelength λ. The depth inside the resin is z.

The molar absorptivity of the resin, ε, depends upon the source wavelength λ. The depth inside the resin is z. The kinetic equation of the initiator can then be given as,𝑑[𝐼𝑛]𝑑𝑡= −𝐾𝑑[𝐼𝑛](5)The radicals can then react with the double bonds to form longer chains, or form a dead radical or be quenched with dissolved oxygen as depicted by the following three equations. ￀R ∙ +𝐷𝐵𝐾𝑝 → 𝑅 ∙(6)R ∙ +R ∙𝐾𝑡 → 𝑅𝑑𝑒𝑎𝑑(7)R ∙ +𝑂2𝐾𝑡,𝑂2→ 𝑅𝑑𝑒𝑎𝑑(8)Rdead is species produced that destroys one or more radicals. The rate constants used are Kp for propagation of a radical through an acrylate double bond, Kt for termination between two radicals, and Kt,O2 for termination of a radical with an oxygen molecule.

The rate constants used are Kp for propagation of a radical through an acrylate double bond, Kt for termination between two radicals, and Kt,O2 for termination of a radical with an oxygen molecule. R* is non-propagating radicals.The overall rate of initiator decomposition, Ri, is modeled by multiplying the rate constant Kd by the initiator concentration [In]𝑅𝑖 = 𝐾𝑑[𝐼𝑛](9)The kinetic equations for the double bond [DB], live radicals [R∙] and oxygen [O2] can be given as follows:𝑑[R∙]𝑑𝑡= 2𝑘𝑑𝐼(𝑧)[𝐼𝑛] − 2𝑘𝑑[R ∙]2 − 𝑘𝑡,𝑂2[R ∙][𝑂2](10)𝑑[𝐷𝐵]𝑑𝑡= −𝑘𝑝[R ∙][𝐷𝐵](11)𝜕[𝑂2] 𝜕𝑡= −𝑘𝑡,𝑂2[R ∙][𝑂2] + 𝐷𝑂2𝜕2[𝑂2] 𝜕𝑧2(12)The effect of oxygen inhibition and diffusion was explicitly modeled in Eq. 12.

12. Due to the high diffusivity of dissolved oxygen in the photopolymer resin, it was assumed that the oxygen would primarily diffuse from uncured top layers of the sample chamber down to the curing front, competing with double bonds for radicals and significantly slowing down the rate at which the double bonds are converted, thus increasing the gel time. This equation was modified to account for oxygen diffusion in two dimensions as shown in Eq. 13.𝜕[𝑂2] 𝜕𝑡= −𝑘𝑡,𝑂2[R ∙][𝑂2] + 𝐷𝑂2𝜕2[𝑂2] 𝜕𝑥2 + 𝐷𝑂2𝜕2[𝑂2] 𝜕𝑧2(13)The researchers estimated the rate constants, Kt, Kp & Kt,O2 by fitting the simulation results with the experimental data from Fourier-transform infrared (FTIR) experiments[9]. They had suggested that the individual rate constants are not unique and may vary.

They had suggested that the individual rate constants are not unique and may vary. Since the FTIR experiments were conducted at 100 times the intensity of the light used in the ECPL system, it is possible that the effect of oxygen inhibition and diffusion was not adequately captured using the presented rate constants. Hence, these constants were varied to suit the ECPL experimental conditions. Constant rates used for the current system will be specified in the section 5.3.3. Process Plan FormulationThe existing process planning methods for ECPL system are discussed in Jariwala et al.[9] In this work, process planning is used to estimate the exposure times and corresponding bitmaps for the desired part shape and dimensions. MATLAB and COMSOL software are used in the process planning. Part shape and dimensions are provided as inputs. The process planner first estimates the bitmap and exposure time for the first layer.

MATLAB and COMSOL software are used in the process planning. Part shape and dimensions are provided as inputs. The process planner first estimates the bitmap and exposure time for the first layer. Then, the program will connect with COMSOL to modify the time by simulating the effects such as oxygen inhibition and diffusion. After that it will calculate the exposure time for next layers in the same method. The flowchart for estimating the first set of process inputs and subsequent set of process inputs are shown in Figures 4 and 5, respectively. With reference to Figure 4, the first bitmap and exposure time were estimated by optimizing the cured part edge. It is to be noted that the primary function of the first bitmap is not to cure the entire part geometry, but to cure the base and the corresponding edge of the desired part.

It is to be noted that the primary function of the first bitmap is not to cure the entire part geometry, but to cure the base and the corresponding edge of the desired part. For the subsequent layer process shown in Figure 5, the inputs are the bitmap for the first “layer” and the corresponding exposure time and the desired cured part geometry. The dotted lines show the simultaneous dependence of the optimization module on the system module and the material module. The dashed lines show the iterative nature of the loop and the process stops when the entire desired part geometry is cured completely (within a pre-set threshold).4. Revised Irradiance ModelThe existing irradiance model has been demonstrated to be able to predict the exposure profile on the substrate level very well. Experimental validation was also reported in Jariwala et al [8]. However, the experimental validation was provided by projecting a large pixel size bitmap with 5, 10 and 30 seconds exposure time.

Experimental validation was also reported in Jariwala et al [8]. However, the experimental validation was provided by projecting a large pixel size bitmap with 5, 10 and 30 seconds exposure time. As Figure 6 shows, the top region of the irradianceFigure 4: Flowchart for estimating first set of process inputs [9]Figure 5: Flowchart for estimating subsequent set of process inputs [9]model has much more variation than the bottom part, which means that for small amount of curing time, the irradiance model can not estimate the exposure profile to a high degree of accuracy. It will work for cylindrical parts which have no limitations in curvature. However, forspherical lens and aspheric lens, the curing time for each “layer” will be less than 2 seconds, which will induce more difference between the simulation results and experimental results. Figure 7 is an aspheric lens example of the experimental results.

Figure 7 is an aspheric lens example of the experimental results. The figure shows that the dimensional error between simulation and the decreasing experiment desired height which is also related to the exposure time. Hence, the irradiance model needs to be improved to fit the experimental data. Figure 8 shows the comparison of revised irradiance model and previous irradiance model. The blue lines show the old irradiance model and the red lines indicate the new irradiance model.Figure 6: cropped view of the top region of Figure 3increases withFigure 7: Dimensional error vs. sample height5. Revised Rate ConstantIn our previous work, significant differences were observed between simulation results and fabricated parts, as shown in Figure 9. Cross-sectional profiles of cured parts as obtained from 3D laser confocal microscopy are plotted along with simulated part profiles. All samples were exposed at the same intensity for an exposure time of 30s with different pixel size.

All samples were exposed at the same intensity for an exposure time of 30s with different pixel size. Results suggest that parts with large diameters can reach the same height for the same amount of exposure time, but for small diameter parts, the reached heights decrease with decreasing diameter. Simulation results cannot accurately predict the case.With a new UV source, the UV light intensity was changed from 140 to 210 W/m2. Due to this change, new rate constants for propagation and termination were fit to data measured from part fabrication experiments. The physical parameters, experimental conditions and the fitted rate constants in the study are presented in Table 1.

The physical parameters, experimental conditions and the fitted rate constants in the study are presented in Table 1. In addition to the change in UV intensity, itwas found that the intensity was not stable in the lab, so the light intensity is measured every time before doing experiments.Evan after this work, it was observed that the rate constants shown in Table 1 only worked for the large pixel bitmaps (greater than approximate 200 m in diameter). Therefore, to fabricate accurate small parts, process plans have to be manually modified. The process is well understood enough that such manual modifications do enable simulation results to match the experimental results.Figure 8: Examples of revised irradiance model6. Effect of Post ProcessingThe post-process is referred as washing process in the ECPL system. It is used to transform the cured part into a finished product.

Effect of Post ProcessingThe post-process is referred as washing process in the ECPL system. It is used to transform the cured part into a finished product. After the cured part is fabricated, the cured sample is still submerged to the uncured resin in the resin chamber. The uncured resin needs to be removed by gently washing. An inappropriate washing method can lead to damage of the cured gel boundary and also the surface of the sample. The washing method developed in the ECPL system is to use a solution of water and Triton-X to gently remove all traces of uncured monomer, and then apply water to clear the surface. Various solvents such as acetonitrile, hexane, ethyl acetate and a solution of water and soap were also tried. The solution of water and Triton-X (soap solution) is found to be the most effective and repeatable way to clean the samples.

The solution of water and Triton-X (soap solution) is found to be the most effective and repeatable way to clean the samples. AsFigure 10 shows, a good washing method can provide a clear, smooth and clean surface, but for bad washing methods, the surface is very rough and looks like a moon face. However, the current washing method will still wash off partial cured resin. The outside layer of the sample will be sticky and easily to be removed especially for smaller samples based on the experiments.

Therefore, in order to reduce the effect of post processing, a dimensional adjustment will be applied before inputting the design specification into the process planning program.Table 1: Physical parameters used in this studyParameterValueUnitsSourceQuantum efficiency of radical, Ф0.6-[17]Molar absorptivity of photons at 365nm wavelength, ε15m2/mol[17,18]UV light Intensity, I0210W/m2experimentalMolecular weight of Monomer, TMPTA296g/molSartomerMolecular weight of Photoinitiator, DMPA256g/molCibaRate constant for propagation reaction, Kp1.66m3/mol-sModified from [15]Rate constant for termination reaction, Kt1.31m3/mol-sModified from [15]Rate constant for termination via oxygen quenching, Kt,O2125m3/mol-sModified from [15]Diffusion coefficient of Oxygen, DO21.00E-10m2/s[19]Initial concentraton of Oxygen, [O2]01.05mol/m3[20]Figure 9: Different pixels with the same amount of exposure time [12]Figure 10: Sample surface with good post processing (right) and bad post processing method (left)7.

Experimental ValidationA set of 42 different part shapes and sizes were investigated in order to test the revised process planning method. This set contains both cylindrical and aspheric parts with diameters ranging from 200 to 590 m and heights ranging from 60 to about 122 m. In Section 7.2, one specific case of an aspheric lens is investigated in detail.7.1 Set of Cylindrical and Aspheric PartsIn order to validate the process planning method with revised irradiance model, revised rate constant and consideration of the effect of the post processing, a large number of cylindrical and aspheric parts were fabricated and measured. Matlab scripts were written to encode the revised process planning algorithm. The bitmaps and exposure time will be obtained after entering the desired geometry. By projecting the bitmaps through the ECPL system, samples were fabricated, then washed and measured using the LEXT 3D confocal microscope.

By projecting the bitmaps through the ECPL system, samples were fabricated, then washed and measured using the LEXT 3D confocal microscope. The experiments and results are presented in Appendix A. In our case, the geometry height and diameter were mainly focused. The desired part height and width, experimental height and width, error between the simulation and experiment in both height and diameter and the error percentage were listed in the table. One sample from the table will be investigated in the following section.7.2 Test CaseThe desired geometry is an aspheric lens with 200 micrometers in diameter and 65 micrometers in height. The conic constant is -1. After running the process planner on this geometry, the bitmaps and exposure times were computed, as shown in Figure 11.The experimental results are shown in Figure 12. A 20X microscope objective was used to obtain the dimensional height and diameter, but measured points along the profile had significant noise.

A 20X microscope objective was used to obtain the dimensional height and diameter, but measured points along the profile had significant noise. Therefore, the 50X microscope objective was used to get the geometry profiles. The figure shows the comparison of desired geometry and experimental result in half-sectional view.Figure 11: Bitmaps used in the test caseFigure 12: Comparison of the half-sectional profiles of the cured part sample and the desired part geometryIt was observed that the experimental result matched the desired geometry very well. The cut-off edge error, between 90 and 100 m in the graph, was likely caused by microscope measurement problems associated with steep part geometry.

The cut-off edge error, between 90 and 100 m in the graph, was likely caused by microscope measurement problems associated with steep part geometry. We can see the edge clearly when the other objective was used.Overall, from the entire table in Appendix A, compared to the previous published results [9], it can be concluded that the current revised processing method with the consideration of the oxygen inhibition and post processing yields much better results. The deviation of part height between the simulation and experiment is less than 10% and the deviation in the diameter is less than 5% which was expected.8. ConclusionThis paper presented a revised process planning method for the ECPL mask-projection stereolithography process with the consideration of oxygen inhibition and post processing. The rate constants of the polymerization model and irradiance model were modified according to the experimental data.

The rate constants of the polymerization model and irradiance model were modified according to the experimental data. By using COMSOL and MATLAB software to simulate the ECPL process, the bitmaps and exposure time which are needed to cure the desired geometry were obtained. As all the experimental result demonstrate, the simulated height and diameter can predict the experimental result fairly well. The error is less than 10% in the height and 5% in the diameter.In order to reduce the error percentage to our demands: less than 5% in both dimension, future improvement need to be investigated. Due to the limitation of the COMSOL, neural network may be used to set up the statistic model to get better results.AcknowledgmentThis material is based upon work supported by the National Science Foundation under Grant No. CMMI-1234561.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEFFECTS OF PARTICLE SIZE DISTRIBUTION ON SURFACE FINISH OF SELECTIVE LASER MELTING PARTSJ H Lim1, N A Khan11 Advanced Remanufacturing and Technology Centre, 3 Cleantech Loop, Singapore 637143AbstractMetal parts produced by Selective Laser Melting (SLM) usually exhibit poor surface finish compared to conventional manufacturing processes. There is a growing need for parts to have good surface quality in the as-built condition to minimise post-processing costs and reduce lead time. There are many studies done on the effects of processing parameters on surface finish but very little on the influence of powder characteristics. This study aims to investigate the effects of Particle Size Distribution (PSD) on surface finish of AM parts by printing coupons with Inconel 625 powders of varying PSD.

This study aims to investigate the effects of Particle Size Distribution (PSD) on surface finish of AM parts by printing coupons with Inconel 625 powders of varying PSD. It was found that roughness of internal surfaces was mainly caused by the presence of partially sintered particles. Whilst a smaller particle mean size and wider particle size range are preferred for better surface finish, a powder that is too fine may result in poor flowability affecting its processability in terms of layering and powder bed quality.IntroductionLaser Powder Bed Fusion (L-PBF) is an Additive Manufacturing (AM) process in which powder materials fuse together layer by layer to form a three dimensional (3D) part based on a computer aided design (CAD) model [1]. For fabrication of metal components, this process is often referred to as Selective Laser Melting (SLM) or Direct Metal Laser Sintering (DMLS).

For fabrication of metal components, this process is often referred to as Selective Laser Melting (SLM) or Direct Metal Laser Sintering (DMLS). During the fabrication process, a thin layer of metal powder material is deposited onto a build platform with a roller or coater blade. A focused laser beam scans the powder bed based on the cross-section of the three dimensional CAD model to selectively fuse the layer of powder. The build platform is then lowered by a fixed layer thickness before a new layer of powder is being deposited on the powder bed. The laser beam scans the powder bed again and the process of joining materials layer by layer repeats itself until the whole three dimensional part is produced.It is widely known that AM is capable of producing highly complex parts with faster processing time [2]. However, parts produced by L-PBF usually exhibit poor surface finish compared to conventional manufacturing processes.

However, parts produced by L-PBF usually exhibit poor surface finish compared to conventional manufacturing processes. Post-processing techniques, such as shot peening, machining and polishing, are often required to achieve desired surface finish [3]. These post-processes can be costly and challenging especially for parts with small or intricate internal features such as conformal cooling channels used for tooling. For AM to reach a state of maturity for industrial applications, there is a growing need for parts to have good surface quality in the as-built condition to minimise post-processing cost and reduce lead time [4].Surface quality of as-built part is dependent not only on part geometry and process parameters [8-10] but can also be influenced by feedstock powder characteristics such as particle size distribution (PSD) [5-7]. PSD refers to the amount of large and small particles in the powder.

PSD refers to the amount of large and small particles in the powder. Energy supplied by the laser may be insufficient to fully melt the large particles and the partially melted particles may result in surface undulation causing an increase in surface roughness.1990However, if sufficient small particles are present, it may help to reduce this effect as small particles can fill into the voids between large particles and are easily melted. Hence, a powder with lower particle mean size and wider particle size range tend to produce parts with better surface finish.Surface quality can also be affected by the effective layer thickness (teff) [11-13], which is the actual powder layer thickness deposited on the part after each scanned layer and is always higher than the theoretical layer thickness (t0). For metal powders used in SLM, the powder layer density typically lies between 40% to 60%.

For metal powders used in SLM, the powder layer density typically lies between 40% to 60%. Assuming that the first layer of powder is deposited at a layer thickness of 30μm and has a powder layer density of 60%, the laser scans that layer to a bulk density of 99%. This leads to a shrinkage in bulk layer thickness to approximately 18.2μm, a difference of 11.8μm. The build platform then lowers by the theoretical layer thickness of 30μm for the next layer of powder to be deposited. This results in an effective layer thickness of 41.8μm as illustrated in Figure 1.1) First layer deposition2) Laser scanning3) Next layer depositionCoaterLaserCoater30µm {18.2µm {Build PlatformFigure 1: Development of effective layer thickness in SLMAssuming a constant mass before and after laser scanning and the shrinkage occurs in the Z-direction, teff can be calculated based on the powder layer density and part density using the following equations.

For example, if teff < D90 of the powder, particles that are larger than teff will not be deposited in the powder layer as they will be pushed across the build surface by the coater blade. The effect of large particles will be less significant and the presence of small particles will have a greater effect in improving surface finish. However, if teff > D90, large particles can be deposited in the powder layer so surface roughness will be attributed to the size and quantity of the large particles present. The effect of small particles1991become less significant in improving surface finish. This means that the higher the effective layer thickness, the greater the influence of large particles and the lower the influence of small particles on surface finish.Although it has been suggested the use of fine particles for better surface finish [11-13], most studies are limited to external surfaces of simple geometries i.e. top and side surfaces of cubes and walls.

top and side surfaces of cubes and walls. The study on the influence of powder characteristics on surface quality of internal surfaces would be of greater interest and importance because these surfaces are hard to reach and more challenging to post-process. This study aims to investigate the effects of PSD on surface finish of internal channels comprising of inclined and overhanging surfaces.Methods and MaterialsGas atomised IN625 powders of three different PSDs were requested from the same production batch with specifications shown in Table 1.

The powders were sampled following ASTM B215-15 scoop method and characterised based on the techniques and standards shown in Table 2.Table 1: IN625 powder specificationsPowderSpecification RemarksA15 – 45 μmTypical size range for L-PBFB0 – 45 μmIncrease in amount of fine particles / wider particle size rangeC0 – 32 μmDecrease in particle mean sizeTable 2: Powder characterisation techniques and standardsPowder CharacterisationTechniques / EquipmentReferenceParticle Size Distribution (PSD)Static Image Analysis (Morphologi G3)ISO 13322-1:2014True DensityHelium PycnometryASTM B923-16Hall Flowmeter FunnelASTM B212-17Apparent DensityCarney FunnelASTM B417-13Tap DensityTap DensityASTM B527-15Hall Flow RateASTM B213-17FlowabilityCarney Flow RateASTM B964-16Layering assessment and coupon fabrication were performed for each powder in virgin condition at a layer thickness of 40μm.

All coupons were processed using standard OEM process parameters and post-processing procedures i.e. heat treatment and EDM wire-cutting for part removal. Each build consisted of coupons with internal channels of 5mm diameter printed vertically and horizontally, as well as density cubes, tensile bars and powder capsules. The build layout is shown in Figure 2.1992(a) Build design(b) As-built partsFigure 2: Build layout of couponsThe powder capsules were used to collect the powder in the build chamber during the fabrication process.

The parts were characterised based on the techniques and standards as shown in Table 3.Table 3: Part characterisation techniques and standardsPart CharacterisationTechniques / EquipmentReferenceDensityArchimedes PrincipleASTM B311-17HardnessMicro Vickers Hardness TestingASTM E384-17Tensile StrengthTensile TestingASTM E8-16aThe coupons with vertical and horizontal internal channels were sectioned using EDM wire-cutting for inspection of the internal surfaces. Surface quality was characterised with areal surface texture parameters i.e. Sa and Sz measured using the Taylor Hobson Talysurf CCI-HD, a non-contact 3D optical profiler. Compared to conventional line measurements that provide 2D roughness profiles, areal analysis provides 3D surface topography information which can be more comprehensive and useful for characterising highly irregular AM surfaces. A 20X magnification lens was used to give a spot area of 0.83mm x 0.83mm for each measurement.

A 20X magnification lens was used to give a spot area of 0.83mm x 0.83mm for each measurement. The features measured are shown in Figure 3. Five coupons were fabricated for each feature and the measurements were taken at three locations for each coupon. An overall average was used for analysis.1993(a) External Surfaces(b) Internal ChannelsVertical SurfaceSide SurfaceTop Surface (Up-skin)Vertical CouponHorizontal CouponInclined Surface (Up-skin)Overhang Surface (Down-skin)Figure 3: Features measured for surface quality evaluationResults and DiscussionPowder CharacterisationThe static image analysis was done using the Malvern Morphologi G3, which captures the 2D image of the particle and converts it into a circle of equivalent area reported as Circle Equivalent (CE) diameter.

It provides a number-based distribution (n) which is more sensitive to smaller particles and is more effective in characterising metal powders in L-PBF containing relatively fine particles. The results are shown in Table 4. Span is a measure of distribution width defined as:Span =(cid:3005)(cid:2877)(cid:2868)(cid:2879)(cid:3005)(cid:2869)(cid:2868)(cid:3005)(cid:2873)(cid:2868)(4)A higher span represents a wider particle size distribution. Results showed that Powder C had the lowest particle mean size while Powder B had the highest span, suggesting a wider particle size range due to the presence of both small and large particles.Table 4: PSD results from number-based distribution C PowderABD[n,0.1] (μm)19.7015.2113.25D[n,0.5] (μm)27.4225.5723.01D[n,0.9] (μm)40.7436.5931.76Mean (μm)29.2726.1423.49Span (n)0.770.840.80True, apparent and tap densities were measured using samples from the feedstock powder before each build.

The powder layer density was calculated from equation (3) using the powder collected in the capsules after each build. True density, which is the absolute density of the powder material IN625, was measured to be 8.55g/cm3 using helium pycnometry.

True density, which is the absolute density of the powder material IN625, was measured to be 8.55g/cm3 using helium pycnometry. This value was considered as 100% density for the computation of relative densities as shown in Table 5.1994Table 5: Density (g/cm³) and relative density (%) for different powdersPowderTrue Density (g/cm³)Apparent Density (g/cm³)Tap Density (g/cm³)Powder Layer Density (g/cm³)A8.55 ± 0.01 (100%)3.89 ± 0.02 (45.50%)4.79 ± 0.02 (56.02%)4.47 ± 0.02 (52.28%)B8.55 ± 0.01 (100%)4.02 ± 0.02 (46.90%)5.01 ± 0.04 (58.60%)4.54 ± 0.01 (53.10%)C8.55 ± 0.00 (100%)3.95 ± 0.02 (46.20%)4.96 ± 0.02 (58.01%)4.51 ± 0.01 (52.75%)Apparent density provides an indication to a powder’s natural packing ability upon free flowing through a nozzle while tap density measures the powder’s ability to re-arrange itself upon mechanical tapping.

However, it may not be accurate representations of powder behaviour in L-PBF where the powder is deposited on the build platform by a coater blade layer by layer. The powder layer density measurement would provide a better estimate of powder density, hence would be used to calculate the effective layer thickness for each powder.From the results, Powder B had the highest apparent and tap densities followed by C and A. Powder layer density displayed a similar trend with values in between apparent and tap density. The difference in powder densities could be correlated to the particle size range where Powder B was shown to have the highest span followed by C and A. The higher the span, the higher the powder densities.

The higher the span, the higher the powder densities. This is because with a wider particle size range, small particles can fill into the voids between large particles allowing them to be packed closer to each other.Powder flowability is influenced by inter-particle interactions such as frictional forces, mechanical interlocking, and adhesion due to moisture ingress. These factors are in turn affected by PSD, morphology, and surface chemistry of the powder. Powder flowability is conventionally represented by Hall and Carney flow rates, which measures the time taken for a fix mass of powder to flow through a calibrated funnel. The longer the time taken, the poorer the flowability. Another measure of flowability is the Hausner Ratio (HR), a parameter correlating powder density and flowability. The higher the HR, the poorer the flowability.

All three powders flowed through the Carney flow funnel and the results confirmed that Powder A had the highest flow rate, followed by B and C. The poor flowability of Powder B and C could be due to the presence of small particles resulting in a higher surface area per unit mass for frictional forces to act between the particles and increasing inter-particle cohesion. In general, the smaller the particle mean size, the poorer the flowability. The Hausner Ratio results followed a similar trend and further supports the observation.19961 9 9 7a n d p o w d e r d e n s i t i e ss h o w e d n o c l e a rc o r r e l a t i o n t ot h ef i n a l p a r t d e n s i t y a n d h a r d n e s s v a l u e s .p l a n e s o ft h ec u b e s .T a b l e 7 s h o w st h e d e n s i t y a n d h a r d n e s st e s t i n g r e s u l t s .T h ec h a n g ei n P S Da t8.

The results for Ultimate Tensile Strength (UTS), 0.2% Yield Strength (0.2%YS) and Failure Strain are shown in Table 8. Based on the results, there was negligible difference in tensile properties between coupons fabricated by the three powders. The change in PSD did not show any significant effect on the static mechanical properties of the final parts.Table 8: Tensile test results for horizontal tensile specimensPowderUTS (MPa)0.2%YS (MPa)Failure Strain (%)A950 ± 5647 ± 428.7 ± 1.5B953 ± 12650 ± 928.2 ± 0.5C936 ± 8637 ± 928.1 ± 0.9Effective Layer ThicknessAll coupons were fabricated using a theoretical layer thickness (t0) of 40μm. The effective layer thickness (teff) was calculated based on the powder layer density and part density of each build using equations (1) and (2).

For a theoretical layer thickness of t0 = 40μm, the effective layer thickness for all powders reached above 76μm, which was much greater than the size of the powders (teff > D90). Large particles could be deposited in the powder layer so surface roughness would be attributed to the size and quantity of the large particles present and the effect of small particles would become less significant. Since the main difference between Powder A (15 - 45μm) and B (0 - 45μm) was in the amount of small particles present, the improvement in surface finish of parts produced by Powder B compared to A might not be obvious. However, Powder C (0 - 32μm) would be expected to produce parts with a better surface finish as it has a smaller D90 and mean size compared to A and B.The influence of PSD on surface finish might change for a lower effective layer thickness.

For example, if the theoretical layer thickness is t0 = 20μm, the effective layer thickness achieved will only be about 38μm, which is below the D90 of Powder A (teff < D90) but above that of B and C. The same explanation cannot be used for Powder A as particles that are larger than teff will not be deposited in the powder layer. In this case, the effect of large particles present will be less significant in increasing surface roughness.Surface QualityFor external surfaces, the results for top and side surface roughness in terms of Sa and Sz are shown in Figure 7. Results showed that top surfaces exhibit better surface finish than side surfaces.

Results showed that top surfaces exhibit better surface finish than side surfaces. While Powder C produced the lowest average Sa and Sz values for both surfaces, the differences between the results for all three powders were marginal.)m μ ( a S10.00 9.00 8.00 7.00 6.00 5.00 4.00 3.00 2.00 1.00 0.00)m μ (z S100.00 90.00 80.00 70.00 60.00 50.00 40.00 30.00 20.00 10.00 0.00ABCABCTop (Sa)[J1.981.921.85aTop (Sz)24.5021.8520.92Side (Sa)[I6.556.225.86mlSide (Sz)71.9471.2870.42Figure 7: Surface roughness results for external surfacesFigure 8 shows the top and side surface topographies obtained for each powder. Top surfaces roughness was mainly caused by surface undulation formed from the laser scan tracks and the effect of PSD was not obvious. In contrast, side surface roughness was attributed to the presence of partially sintered particles adhering to the surfaces.

In contrast, side surface roughness was attributed to the presence of partially sintered particles adhering to the surfaces. Naturally, the smaller the size of these particles, the lower the roughness values.1999p o Te d i S~m m- = - ~mm- - - ~~mm - = -~ .A~0m - = -~~ID0 ~ - = -~B0~200u300400wS000 - -- ~ • - w - = - ID=~~~m - =~ID~m m- = -- ~ • - w - -- ~ 0uwm~0~Figure 8: Surface topography for top and side surfacesm - = ~~ ~CIDuw= - - ~µm--0~~w •ID--w0For the internal channels, vertical surface refers to the internal surface of the vertical channel while inclined and overhang surfaces refer to the up-skin and down-skin of the horizontal channel. The surface roughness results in terms of Sa and Sz are shown Figure 9. Results showed that overhang surfaces exhibit the poorest surface finish with the highest Sa and Sz values.

The surface roughness results in terms of Sa and Sz are shown Figure 9. Results showed that overhang surfaces exhibit the poorest surface finish with the highest Sa and Sz values. Inclined surfaces had lower Sa values but higher Sz values compared to vertical surfaces.

Partially sintered particles were seen on all surfaces but in different sizes and quantities for different surfaces.2000ABC2004006008001,1mµm200400600800 µmµm200400600800 µm100200908070100200908070100200l a c i t r e VJOO400500600605040JOJOO400500600605040JOJOO4005006007002070020700800µm100 NM800µm100 NM800µm200400600800µmµm 150200400600800""µm 1502004006008001,1m100100100200200200d e n300400100300400100300400i l c n I5005050060050500600700700700800µm.

NM0800µm0 NM800µm2004006008001,1mµm200100600800 µmµm200400600BOOµm100100100200300200300200g n a h r e v OJOO400500600200100300400500600200100JOO400500600700700700800µm0 NM800µm0 NH800µmFigure 10: Surface topography for vertical, inclined and overhang surfacesVertical surfaces showed similar characteristics to side surfaces in terms of the amount of partially sintered particles. Inclined surfaces seemed to have a lower amount of partially sintered particles but tend to be larger in size and irregularly shaped. This might be the reason for the lower Sa but higher Sz values compared to the vertical surfaces. Surface roughness was also influenced by the staircase effect evident from the dark valleys shown in the surface topography. The staircase effect is unique to inclined surfaces where the higher the layer thickness, the higher the surface roughness.

The staircase effect is unique to inclined surfaces where the higher the layer thickness, the higher the surface roughness. Overhang surfaces had the highest amount of partially sintered particles causing highly irregular surface topographies. This is because during the fabrication, the overhang surfaces were in direct contact with the powder bed and not supported by scanned bulk layers or support structures.ConclusionTo study the effects of PSD on surface finish of SLM parts, IN625 powder of three different PSDs were characterised and processed using 40μm layer thickness, standard processing and post processing parameters. Powder B had the widest particle size range, resulting in the highest powder layer density. Powder C had the smallest particle mean size, resulting in high cohesiveness and poor flowability, which was also evident from the irregularities i.e. scratch lines seen in the first few layers of powder deposition.

scratch lines seen in the first few layers of powder deposition. The change in2001µm9080706050••JO2010. NM0µm ISO100500 NMµm3002001000 NHPSD showed no significant effect on the final part density, hardness and static mechanical properties.The teff for t0 = 40 μm was shown to be greater than the particle size of all powders, suggesting that surface roughness would be attributed to the size and quantity of the large particles present and the effect of small particles would become less significant. Results showed that Powder C produced surfaces with the lowest average roughness values, suggesting that a lower particle mean size led to small improvements in surface finish for both external and internal surfaces. From the surface topographies obtained, it can be seen that the roughness of internal surfaces was mainly attributed to the size and quantities of partially sintered particles.

From the surface topographies obtained, it can be seen that the roughness of internal surfaces was mainly attributed to the size and quantities of partially sintered particles. This means that particle mean size will have a direct influence on surface roughness as the larger the size of the partially sintered particles, the higher the roughness values. teff also plays a part as it determines the amount of large particles that can be deposited in the powder layer. In general, a smaller particle mean size and wider particle size range would be preferred for better surface finish, but a powder that is too fine may result in poor flowability affecting its processability in terms of layering and powder bed quality.Powder behaviour can differ across AM platforms with different layering mechanism. This study was conducted using a stiff coater blade made of steel.

This study was conducted using a stiff coater blade made of steel. It can be extended to evaluate the effects of powder characteristics on surface finish using coater blades of different material, geometry and stiffness, or with a roller coating mechanism. Surface finish can also be affected by the chemical properties of powder. Further research can include analysis on surface chemistry, moisture content, oxide layer formation, or melt pool characteristics and their effects on surface finish.2002References[1] M. L. Vlasea, B. Lane, F. Lopez, S. Mekhontsev, A. Donmez, 26th Annu. Int. Solid Free.Fabr. Symp. 2015, 527.[2] N. Guo and M. C. Leu, “Additive manufacturing: Technology, applications and researchneeds,” Front. Mech. Eng., 2013, vol. 8, no. 3, pp. 215–243.[3] Mumtaz, K. and Hopkinson, N., Top surface and side roughness of Inconel 625 parts processed using selective laser melting. Rapid Prototyping Journal, 2009, 15(2), 96–103.[4] D.L. Bourell, H.L. Marcus, J.W.

fects of Post Processing Heat Treatments on the Bond Quality and Mechanical Strength of Ti/Al3003 Dual Materials Fabricated using Ultrasonic Consolidation J.O. Obielodan, B.E. Stucker, Utah State University.AbstractThe interface between layers in ultrasonically fabricated parts is often poor for desirablematerial combinations, resulting in relatively low bond strength. This makes these fabricationsunsuitable for structural applications. This work discusses a study of the effects of postprocessing heat treatment of ultrasonically consolidated titanium and aluminum dual-materialspecimens. The shear strengths of as-deposited specimens as well as heat treated ones weretested. The results show that there is significant improvement of the strengths of post processedspecimens over the as-fabricated ones.

The results show that there is significant improvement of the strengths of post processedspecimens over the as-fabricated ones. The improvement is as a result of interactions of the basematerials across the interfacial boundaries at elevated temperatures, leading to stronger bonds.The study highlights the role of post processing for improving the mechanical properties ofultrasonically consolidated structures.1.IntroductionUltrasonic consolidation is a solid-state fabrication process that combines ultrasonicmetal welding and layered manufacturing techniques to produce three-dimensional freeformobjects. The process uses the power of high frequency ultrasonic vibration at low amplitude tobond thin foils of materials to form solid objects. It combines normal and oscillating shear forceson mating foils and the resulting friction forces between the materials to fracture and displacesurface oxides from the materials.

It combines normal and oscillating shear forceson mating foils and the resulting friction forces between the materials to fracture and displacesurface oxides from the materials. The exposed atomically cleaned surfaces are then brought intodirect contact under modest pressure and temperatures that are less than half of the melting pointof the materials. The materials are thus metallurgically bonded [1]. Fractured oxides and surfaceimpurities in the materials are distributed in the bond zone. The process combines the layer-by-layer addition of foils with contour milling using the integrated 3-axis CNC machining facilitiesto produce desired component geometry.

The process combines the layer-by-layer addition of foils with contour milling using the integrated 3-axis CNC machining facilitiesto produce desired component geometry. It is therefore both an additive and subtractive process.Notable advantages of the solid state nature of the UC process are discussed by White [1].The UC machine consists of a welding horn, also known as a sonotrode, that exerts anormal force and the oscillatory high-frequency vibration on the materials to be welded. Weldingtakes plate on a substrate fixed on a heated plate or anvil. The UC machine is designed forautomatic foil material feeding, but materials can also be fed manually. Figure 1 shows theschematic view of the ultrasonic consolidation process. The primary process parameters arevibration amplitude, temperature, welding speed, and normal force [2].

Figure 1 shows theschematic view of the ultrasonic consolidation process. The primary process parameters arevibration amplitude, temperature, welding speed, and normal force [2]. Other parameters that canaffect weld qualities include sonotrode roughness, materials surface finish [3], and side-by-sidefoil positioning accuracy with respect to the automated material feed system [4].Figure 1: Schematic of UC processUltrasonic consolidation technique is applicable for rapid tooling for injection molding,extrusion, vacuum forming tools and others. It is also used for fabricating tools with conformalcooling channels [1].

It is also used for fabricating tools with conformalcooling channels [1]. Previous works have demonstrated other potential applications for UC.These include honeycomb structures [5], embedding shape memory alloy (SMA) fibers andsilicon carbide fibers in aluminum matrices [6-9], and embedded electronics structures [10].While the process has been widely used for single material fabrications using aluminum alloys, afew researchers have demonstrated its capabilities for multiple material fabrications. The multi-material capabilities of UC was demonstrated by Janaki Ram et al. [11] in their work in whichcopper, brass, nickel, inconel 600, AISI 347 stainless steel, stainless steel AISI 304 wire mesh,MetPreg, and aluminum alloy 2024 were individually welded to aluminum 3003 H18 materials.Domack et al., [12], demonstrated the capability of UC for graded materials compositionfabrications using titanium and nickel alloys. Obielodan et al.

Obielodan et al. [13] have also demonstrated UCmulti-material capabilities by welding different combinations of molybdenum, tantalum,titanium, copper, silver, nickel aluminum alloys 1100, 3003, 6061 and boron powder.The bond strengths of UC fabricated structures is a major concern in attempts to use themfor mechanically stressed structural applications. Some earlier works have determined inter-layerbond related mechanical properties of ultrasonically consolidated structures. Kong et al. [14]determined the peel strength of Al alloy 6061 and Tuttle R.B. [15] determined the peel strengthbetween stainless steel 316L foils welded on a stainless steel plate. Also, Yanzhe Yang et al. [8]determined the peel strength of aluminum alloy 3003-H18 with embedded silicon carbide fibersin aluminum matrix, and Obielodan et al.

Also, Yanzhe Yang et al. [8]determined the peel strength of aluminum alloy 3003-H18 with embedded silicon carbide fibersin aluminum matrix, and Obielodan et al. [4] determined the conditions for optimum transversetensile strength of ultrasonically consolidated structures made with automatically fed aluminumalloy 3003-H18 foils. While most of the previous works were on similar materials, there is agrowing interest in multi-material structures fabrication by ultrasonic consolidation. A majorlimiting factor is that many materials are not easily joined at the current limits of operatingparameters of available UC machines. Some material combinations will require higher values ofparameters that are beyond the upper bounds of the machines currently available. The presentwork seeks to apply post process heat treatment as a way to improve the bond strength betweenultrasonically consolidated titanium and aluminum alloy 3003.

The presentwork seeks to apply post process heat treatment as a way to improve the bond strength betweenultrasonically consolidated titanium and aluminum alloy 3003. Although the low temperatureoperating conditions of UC is a major advantage as stated earlier, it is postulated that by applyingpost processing treatments such as heat and pressure at optimized levels, significant inter-layermaterial diffusion across the consolidated foil interface will be achieved resulting in better bondstrengths. It is a well known fact that high temperature treatment of some dual materials lead tothe formation of brittle intermetallic phases; this study seeks to avoid or minimize theirformation.The UC post process heat treatment in this work is limited to elevating UC fabricatedspecimens to higher temperatures without pressure application. The improved strengthadvantages derivable from the post process heat treatments will most probably offset theadditional costs incurred.

The improved strengthadvantages derivable from the post process heat treatments will most probably offset theadditional costs incurred. The study seeks to create a synergy by combining the freeformfabrication capabilities of UC additive manufacturing with the inherent advantages of diffusionbonding. Improved diffusion based interlayer bond strengths and achievable complex geometrieswill expand the scope of applications for UC structures. The success of this work with titaniumand aluminum 3003 will lay the ground work for further investigations and eventualcommercialization for aerospace applications and in other industries.2.Experimental Work2.1 Material PreparationA Solidica FormationTM machine was used for all the UC fabrications in this work. Thematerials used were annealed commercial pure (CP) Titanium foils of 75 microns thickness andAluminum alloy 3003-H18 of 150 microns thickness with chemical compositions shown inTable 1.

Thematerials used were annealed commercial pure (CP) Titanium foils of 75 microns thickness andAluminum alloy 3003-H18 of 150 microns thickness with chemical compositions shown inTable 1. Deposits were made on Aluminum 3003-H14 substrates. Aluminum substrates and foilswere procured from Solidica. The titanium and aluminum foils were welded on the substrate inalternate layers for four layers, with titanium foils being the first to be welded onto the substrate.Different welding parameters were applied for the two materials as listed in Table 2.

Thewelding temperature was maintained at 300oF (150oC) for the two materials.Table 1: Chemical compositions of the materials used and their form MaterialNominalTensileShearCompositionStrength (MPa) Strength (MPa)CP Titanium Ti-0.6Fe-0.38Mn430380Thickness (µm) 75Al 3003-H18 Al-1.2Mn-0.12Cu200110150Table 2: Ultrasonic welding parameters used for the different materials MaterialAmplitude (µm) 28Speed Normal Force (mm/s) (N) (oF) 300 10.58TemperatureCP Titanium2000Al 3003-H18 1623.701750300Titanium foils were manually placed on the substrate for welding while the aluminumfoils were automatically fed. Other layers of materials were welded on top of the first four layersin order to strengthen the welded foils and avoid tensile failure outside the joints as reported fortwo single aluminum 3003 foils tested for lap shear tests by Kong et al.

[2].The shear testspecimens, described in Fig.2 were designed to fail in shear along lapping surface between thefirst titanium foil and the aluminum 3003-H14 substrate material. During initial specimen trials,Figure 2: Lap shear strength designslots of 3.2mm width and 3mm depth were machined into the Al 3003 substrate with theintegrated CNC facilities in the UC machine, before the first titanium foil deposit was made. Theinitial slot was to provide a precise separation between the ends of the tensile shear specimens atpredetermined location. It was found that the foils deposited above the slots were not bonded toeach other as there was no support material underneath to resist the normal force of the sonotrodefor proper welding. The titanium foils exposed by the cut slots were thus highly oxidized duringheating in the oven.

The titanium foils exposed by the cut slots were thus highly oxidized duringheating in the oven. These specimens failed prematurely because the tensile loads applied werenot evenly distributed, and the exposed titanium foils were weakened by oxidation. Withsubsequent specimen trials, the slots were not cut into the substrate before the foil welding. Theywere cut using a milling machine after material deposition and post process heat treatment werecompleted.Eighteen specimens were fabricated and randomly grouped into six groups of threespecimens each. Each group was then randomly assigned to a post process annealing at 480oCfor the following length of time: 0 (control), 30, 60, 120, 180 and 270 minutes, after which theywere oven cooled to room temperature. The treatments are respectively labeled A, B, C, D, E andF for the purpose of analysis.

The treatments are respectively labeled A, B, C, D, E andF for the purpose of analysis. Those assigned to zero minutes (treatment A) were not heat treated,and served as the control specimens for comparison with the groups that were subjected to postprocess annealing. A Lindberg BlueM laboratory table top oven without atmospheric control wasused for the post process annealing treatments. The foil lapping surfaces were not exposed to theoven atmospheric conditions, only the edges (which are very small compared to the total surfacearea of the titanium foils) were exposed, and as such minimal lapped surface oxidation couldoccur. The samples were loaded into the oven for annealing without the substrates machined offin order to avoid heat induced distortion in the specimens.

The samples were loaded into the oven for annealing without the substrates machined offin order to avoid heat induced distortion in the specimens. After removal from the oven, thesubstrate materials were machined down to 3mm after which the slot was cut to separate the twoends of the tensile shear specimens, which were then only joined at the overlapping surface ofthe welded foils as shown in Figure 2. Because the preparations for the slot machining weremanually done, the slot cuts resulted in different overlap lengths for the specimens. The resultingoverlap lengths of the specimens ranged from 2.32 to 3.3mm. It is assumed that the differencesdoes not significantly affect the lap shear strength measurements obtained.A 50KN capacity Tinius Olson tensile testing machine was used for the lap tensile sheartest. The specimens were held in the flat grips of the testing machine and pulled in tension at aspeed of 1mm/min until fracture.

The specimens were held in the flat grips of the testing machine and pulled in tension at aspeed of 1mm/min until fracture. The maximum fracture load was divided by the area ofoverlapping surface to arrive at the failure shear stress for each specimen.2.2 MetallographySmall samples cut from the unstrained grip ends of the tensile shear strength specimenswere mounted and polished according to standard metallographic procedures. They were thenetched with Kellers reagent (90ml H2O, 5ml HNO3, 3ml HCl and 2ml HF) and observed in opticalmicroscope (OM) and scanning electron microscope (SEM). Line scan Electro Dispersed X-ray(EDX) analysis across the interface of the titanium and Al 3003-H18 deposits were undertakenfor one representative specimen randomly selected from each of the five specimen groups thatwere annealed.

This was done to verify if post process heat treatment-induced diffusion tookplace (and the distribution of diffusing elements, if it occurred) across the interface of theconsolidated materials. The lack of atomic diffusion across the interface of ultrasonicallyconsolidated dual material foils in previous experiments indicated that it was unnecessary tocarry out EDX analysis on any of the control specimens. It is assumed that detectable diffusionwill not occur at the UC operating parameters for the as-deposited samples.3Results and Discussion3.1 MicroscopyThe optical and SEM micrographs of the specimens are shown in Figures 2-8.

The arrowsin Fig.2a show the titanium foils.CP TitaniumAl 3003Figure 2: OM micrograph of a specimen without post processingCP TitaniumAl 3003a: OM micrograph of a 30 minutes annealed specimenAl 3003CP Titaniumb: SEM micrograph of a 30 minutes annealed specimenFigure 3: Sample OM and SEM micrographs of 30 minutes annealed specimensCP TitaniumAl 3003CP TitaniumFigure 4: OM micrograph of a 60 minutes annealed specimenAl 3003CP TitaniumAl 3003Figure 5: OM micrograph of a 120 minutes annealed specimenCP TitaniumAl 3003Figure 6: OM micrograph of a 180 minutes annealed specimenAl 3003CP TitaniumAl 3003Figure 7: OM micrograph of a 270 minutes annealed specimena: SEM micrograph showing the fracture surface of a 30 minute annealed specimenFractured CP TitaniumAl 3003b: SEM show one of the fracture mode of a 30 minute annealed specimenFigure 8: SEM of the fracture features of the specimensThe micrographs in Figures 2 to 7 shows that the bonds between consolidated titaniumand aluminum alloy foils were generally good.

There are no visible bond defects between thelayers. Figure 8 shows the fracture features of the specimens. The fracture morphology at thelapped surfaces is shown in Fig.8a. Figure 8b shows how in some of the specimens, the firstconsolidated titanium layer that bridge the two halves of the lap shear specimens failed intension. The tensile failure of the first consolidated titanium layer is most probably due to thedelamination of that layer that occur as a result of induced bending moment about the center ofthe lapping surface as the applied tensile shear load reaches a particular level (Fig.9 shows thepicture of the bent profile of a 30 minute annealed specimens). The delamination causesdifferential strain between the separated foil and the undelaminated ones. The delaminated foilfails early because the applied stress soon exceeds its tensile strength as its thickness is relativelysmall.

The delaminated foilfails early because the applied stress soon exceeds its tensile strength as its thickness is relativelysmall. In those cases, the lapping foil fraction is completely separated from the longer sidecausing the immediate consolidated aluminum foil to fracture by shear as the load increases.

In those cases, the lapping foil fraction is completely separated from the longer sidecausing the immediate consolidated aluminum foil to fracture by shear as the load increases. Theoutput data obtained from the lap shear strength tests are shown in Table 3 and Figure 10.Figure 9: A lap specimen with bent lapped sectionTable 3: Lap Shear Strength (MPa) data Samples Annealing Time (minutes)1 44.302 31.4803 37.56Average 37.783069.3462.7386.8272.966051.1743.0670.7354.9912053.8133.5341.9643.1018046.2532.2653.8444.1227053.6169.9650.8758.15Figure 10: Bar chart of the lap shear strengths of the different groups of specimens.SAS 9.1 software was used for the statistical analysis of the data to verify the effect of thepost process annealing as a single factor with six treatment levels on the shear strengths of thespecimens.

The boxplot from the analyses did not show any outlier and the normal quantile plotis close to a straight line. Also, all the tests for normality have high P-values, so the assumptionof approximate normality of the data is satisfied. The results of the analysis also show that thereis a uniform spread of the errors, which means the data satisfies the homoscedasticityassumption. The analysis of variance (ANOVA) in Table 4 with a P-value of 0.0210 shows thatpost process annealing has a significant effect on the lap shear strengths for the different groupsof specimens. A post hoc mean comparison using the Ryan-Einot-Gabriel-Welsch multiple rangetest (REGWQ) method is shown in Table 5. The REGWQ controls the maximum experiment-wise error rate under any complete or partial null hypothesis. The table shows that the specimenswith 30 minutes post process annealing (treatment B) yield the highest mean shear strength of72.963MPa.

The table shows that the specimenswith 30 minutes post process annealing (treatment B) yield the highest mean shear strength of72.963MPa. Since treatment B shares the same REGWQ group A with specimens that wereannealed for 270 minutes (treatment F) with mean shear strength of 58.147MPa and with thoseannealed for 60 minutes (treatment C) with mean shear strength of 54.987MPa, it shows that itsdifferences in result with those two treatments groups are not statistically significant.

It (that is,treatment group B) however, have statistically significant higher mean shear strength thanspecimens that were annealed for 120 minutes (treatment D), 180 minutes (treatment E), and thecontrol specimens (treatment A).Table 4: Result of Analysis of Variance (ANOVA) for the dataSourceDF Sum of SquaresMean SquareF ValuePr > FModel52488.774111497.7548224.100.0210Error121457.120067121.426672Corrected Total173945.894178Table 5: REGWQ post hoc means comparisonMeans with the same letter are not significantly different.REGWQ GroupingMean N treatmentA72.9633 BABA58.1473 FBABA54.9873 CBB44.1173 EBB43.1003 DBB37.7803 AThe result of the line scan EDX analysis across the interface of the consolidated dual materialsreveals that diffusion took place at the 480oC annealing temperature. Figure 11 shows thediffusion trend of aluminum into the titanium foil at the interface.

Figure 11 shows thediffusion trend of aluminum into the titanium foil at the interface. It can be seen that the extent ofdiffusion of aluminum into titanium was different for the different post process annealingdurations. The Figure shows that the specimens that were annealed for 30 minutes had about 13wt.% composition of aluminum at the surface of the titanium foil. At a depth of 0.2 microns intothe titanium surface, the aluminum composition is about 7 wt.%. From the titanium-aluminumphase diagram shown in Fig. 12, alpha titanium (α-Ti) can hold up to 7 wt% aluminum in solidsolution at 480oC, beyond which apha2 titanium (α2-Ti3Al), an intermetallic phase, graduallybegins to precipitate. It therefore means that within a depth of 0.2 microns from the surface ofthe titanium foil, α-Ti and α2-Ti3Al coexist after annealing.

It therefore means that within a depth of 0.2 microns from the surface ofthe titanium foil, α-Ti and α2-Ti3Al coexist after annealing. There is the posibility of thesematerials interfacing with a very narow band of other titanium aluminide intermetallics on thealuminum side of the interface. Figure 11 shows higher weight percent concentrations ofaluminum diffused deeper into titanium foils for longer annealing times than 30 minutes. Usingthe titanium-aluminum phase diagram, it can be seen that thicker layers of titanium aluminideintermetallic phases must have formed at the interface.

The specimen group F (with 270 minutesannealing) recorded the highest depth of high concentration of aluminum diffusion into thetitanium foils.Figure 11: A plot of line scan EDX analysis result showing diffusion trend of aluminum into titanium at the interface at different annealing timesFigure 12: Titanium-aluminum (Ti-Al) phase diagram [16]It can be infered that the group B specimens yielded the highest average tensile shearstrength because of the significant aluminum in solid solution of the alpha titanium at theinterface, which co-existed with the narrowest intermetallic phases that may have formed.Specimen groups C, D and E must have failed at relatively lower shear stresses as a result of thehigher depths of brittle intermetallics occassioned by the higher depth of penetration ofaluminum below the titanium surface.

Because the intermetallics are hard, their thicker layersproduced in the group F specimens must have offered some resistance to early failure therebyyielding relatively higher shear strengths than the C, D and E groups. The shear strengthsrecorded for the thicker intermetallic layer is however not reliable under service conditions. It isalways better to avoid or minimize the formation of the intermetallic phases for reliability underservice conditions. This makes the group B treatment prefered over the other treatments.Although the mechanical properties of the CP titanium used is yet to be experimentallyverified, it is generally known to be superior to those listed for aluminum 3003-H18 in Table 1.It is also known that the shear strengths of the consolidated titanium-alumnium dual materialscan not exceed that for Al 3003-H18 base material of 110MPa.

Despite falling short whencompared to the shear strength of the aluminum alloy 3003-H18 shear strength, the highestaverage of 72.96MPa for the post conslidation annealed titanium/Al 3003-H18 is a significantimprovement over the 37.78MPa average shear strengths obtained for the as-consolidatedspecimens. The effects of annealing must have also contributed to the reduced shear strengthsobtained, as the base materials will undergo softening in the process. It is assumed that if heattreatable aluminum alloys are used, better shear strength values can be obtained. Also, the as-consolidated specimens may not have uniformly perfect bonding, which can affect the ultimatestrength values of the post processed specimens. Application of pressure as another factor isexpected to offer improved results.4ConclusionsThis experiment has shown that post consolidation heat treatment by annealing at 480oChas significant effects on the lap shear strengths of UC specimens.

The best average lap shearstrength of 72.96MPa was obtained for specimens heat treated for 30 minutes compared to37.79MPa for as-consolidated ones. Higher heat treatment times result in the formation of thicklayers of brittle titanium aluminide intermetallic phases that causes less reliability and earlyfailure under service conditions.AcknowledgementThe authors wouldliketotakethe opportunitytothankthe NationalScienceFoundation(under GrantCMMI061457)andthe OfficeofNaval Research (under Grant No. N000140710633) for financially supporting this work. Thetechnical support of Solidica, Inc. is also gratefully acknowledged.References1White D.R., 2003. “Ultrasonic Consolidation of Aluminum Tooling”, Advanced Materials & Processes, Vol.161, No.1, pp.64-65.2Kong C.Y., Soar R.C., Dickens P.M., 2004.

e effects of powder feedstock and process parameters on the material characteristics of Ti6Al4V thin strut features fabricated by laser powder bed fusion additive manufacturingNaresh Koju a, Jonah Hermesb, Sumit Paula, Sayed Ehsan Saghaiana, Li Yangaa Department of Industrial Engineering, University of Louisville, Louisville KY 40292, USA b Department of Mechanical Engineering, University of Louisville, Louisville KY 40292, USAAbstractIn this work, three different types of Ti6Al4V powder feedstock of different particle size ranges (fine, medium, and coarse) were utilized to fabricate thin strut lightweight features using laser powder bed fusion additive manufacturing (L-PBF-AM) using different process parameter settings. Thin strut features of varying dimensions from 0.1mm to 0.5mm were fabricated.

Thin strut features of varying dimensions from 0.1mm to 0.5mm were fabricated. The resulting sample sets allow for the analysis of the compound powder feedstock-process- geometry-material (PPG-M) characteristics for lightweight features fabricated by L-PBF-AM, which have not been previously explored. Various material characteristics were experimentally determined and analyzed, including success rate, geometry quality, porosity, pore size, grain size, and mechanical properties of the lightweight thin strut samples. The results clearly demonstrated the significance of the compound PPG-M relationships for lightweight structures, which calls for further studies to “re-establish” the knowledge base for L-PBF-AM materials at small dimension scales.KeywordsLightweight features, thin struts, powder feedstock, powder bed fusion, porosity, microstructure, strength1.

IntroductionCellular structures are composed of interconnected networks of walls or struts or small unit cells (periodic or random) to achieve unique benefits such as a high strength-to-weight ratio, transfer control, and minimal material enhanced mechanical energy absorption, heat requirements [1, 2]. These unique benefits have attracted their use in different engineering applications such as biomedical implants, aerospace structures, automotive parts, sound insulation and damping, electrodes, heat exchangers, and other general lightweight designs [1-3]. Additive manufacturing (AM) is considered to be uniquely suitable for the realization of complex cellular structure designs due to its freeform geometry generation capabilities. However, one of the limiting factors for the adoption of AM cellular structure is the quality characteristics of the materials [4].

However, one of the limiting factors for the adoption of AM cellular structure is the quality characteristics of the materials [4]. The cellular structure usually requires multiscale analysis involving unit cell macroscale and thin strut/wall microscopic levels. The responses at the microscopic structure level (i.e. cellular walls or struts) are more significantly influenced by the local material and process characteristics, as well as material anisotropy, which in turn affects the overall properties of the cellular structure [5]. Furthermore, simplistic periodic lightweight designs that are generally studied in the literature are actually not ideal for applications, and2120often times optimization of lattice structures with varying wall thickness or strut diameters is needed for realistic designs [6].

Therefore, studies on the material characteristics of these thin features i.e., thin wall and thin strut are of particular interest.Studies on the relationships between the material and mechanical characteristics of the thin features fabricated by laser powder bed fusion (L-PBF) and various design aspects such as feature geometry and the process parameters are available in the literature [7-15]. Current knowledge suggests rather complex correlations between L-PBF process parameters such as laser power and scan speed and the geometrical qualities of the Ti6Al4V micro-pillars (struts) and micro-plates (walls) [9]. The effects of feature orientation on the dimension, mechanical, morphological, and microstructural properties of L-PBF thin struts were also studied [6, 10-13]. Reduced geometrical quality with increased surface roughness and reduced cross-sections circularity was identified at lower-feature orientation angle struts [10, 13]. Murchio et al.

Reduced geometrical quality with increased surface roughness and reduced cross-sections circularity was identified at lower-feature orientation angle struts [10, 13]. Murchio et al. [13] reported that the feature orientation directly affects the surface texture and geometry-induced surface imperfections attributed to different thermal behavior history and staircase effect, and plays a vital role in the thin feature's fatigue life. Vrána et al. [14] developed process window for L-PBF process strut-based cellular structure fabrication using different levels of overlaps of contour scans only to achieve controlled porosity and roughness level based on combination of input energy density.Weißmann et al. [11] performed a comparative study between L-PBF and electron beam powder bed fusion (EB-PBF) of Ti6Al4V struts.

[11] performed a comparative study between L-PBF and electron beam powder bed fusion (EB-PBF) of Ti6Al4V struts. The slenderness ratio, which was defined as the ratio between the effective length and the gyration radius of the fabricated strut, was shown to be correlated to both the compressive strength, equivalent diameter and height of the struts. L-PBF struts showed higher dimensional accuracy, significantly lower roughness Ra and Rz than the EB- PBF struts. In addition, the struts fabricated by the two types of PBF processes exhibit different hardness due to different microstructures induced by their respective thermal histories. Leicht et al. [12] reported that in the PBF process both the part geometry and feature orientation exhibit significant effects on the microstructure and texture of the thin features.

Leicht et al. [12] reported that in the PBF process both the part geometry and feature orientation exhibit significant effects on the microstructure and texture of the thin features. For thin wall features, heterogeneous nucleation takes place from individual powder particles at the boundaries, with subsequent grain growth towards the interior center following the maximum temperature gradient. Since the grain growth near the boundary is inhibited by higher cooling rate, with the increase of thin wall thickness, the amount of large elongated grains with a more pronounced texture oriented parallel to the building direction also increases. A study performed by Wang et al. [15] on stainless steel 316L vertical struts reported no clear trend of columnar grain width along the build direction with variation of the strut diameter.

A study performed by Wang et al. [15] on stainless steel 316L vertical struts reported no clear trend of columnar grain width along the build direction with variation of the strut diameter. However, preferred grain orientation begins to appear with the decrease of strut diameter.Feedstock powder characteristics such as particle size and its interplay with layer thickness plays an important role in the PBF process characteristic and is related to various part quality aspects such as density, surface quality, accuracy, internal defects, static mechanical properties and fatigue properties [16-26]. During the process of PBF, powder flowability/rheology has significant effect on the formation of homogenous powder layer. Generally, the flowability increases with decrease of particle size distribution (PSD) range, improves with coarser particles and decreases with increase in moisture content until saturation point [16, 18, 19, 26]. Balbaa et al.

Balbaa et al. [26] reported the presence of refined cellular microstructure while using the coarse AlSi10Mg powder feedstock, which resulted in higher microhardness with the fabricated parts2121compared to those fabricated using fine powder feedstock. On the other hand, depending on the powder spreading mechanism, a bi-modal PSD distribution with refined powder mixed in coarser powder could benefit the powder bed density and the subsequent part quality, due to the filling of the inter-particle voids among larger particles by the small-size particles [17, 22]. Groarke et al. [27] reported improvement of powder flow with the presence of fine particles in the feedstock. In addition to the powder feedstock PSD, the shape of the particle/ morphology also has significant effects on powder rheological properties [19, 20].

In addition to the powder feedstock PSD, the shape of the particle/ morphology also has significant effects on powder rheological properties [19, 20]. For non-refined powder feedstock, spherical particle shape facilitates the powder packing efficiency and powder flowability compared to non- spherical or irregular-shaped particles [16, 19].In comparison, literature about the effect of powder feedstock on the process characteristics, geometry accuracy and material characteristics of the thin features (< 1mm) is much more limited. The studies on single track powder bed fusion can be useful in understanding of thin feature process characteristics due to their dimensional similarity. A multi-layer single track L- PBF-AM process simulation study conducted by Cao and Guan [28] showed that the surface roughness increases with increase in proportion of large-size particles.

A multi-layer single track L- PBF-AM process simulation study conducted by Cao and Guan [28] showed that the surface roughness increases with increase in proportion of large-size particles. In addition, the presence of higher proportion of small-size particles at smaller layer thickness (30µm) and lower proportion of larger-size particles at larger layer thickness (45µm) led to lower porosity. In general, there still exist significant knowledge gaps in the compound powder feedstock-process- geometry-material (PPG-M) characteristics for lightweight features fabricated using L-PBF. Therefore, in this work, three different powder feedstocks (fine, medium, and coarse) were utilized, and thin struts of varying dimensions were fabricated via various process settings using these powder feedstocks for the characterization of the PPG-M characteristics of these lightweight features.

Various process and material characteristics such as fabrication success rate, geometrical accuracy, porosity, pore size, microstructural grain width and mechanical properties were investigated. Due to the large dimension of the comprehensive PPG-M characteristics, some of the input variables that are expected to be also significant such as feature orientation angle, layer thickness and scan strategy were not included in the experimental design of this study. It was anticipated that the results from this work could provide new qualitative insights into the PPG-M relationships, while more quantitative and comprehensive understanding of the dimension-dependent process characteristics would likely require additional investigation via a combination of both experimentation and process modeling.2.

Experimental methods 2.1.Powder Feedstock characterizationThree powder feedstocks with varying particle size ranges were purchased from Praxair Surface Technologies, Inc., namely Ti-105-1 (Medium powder, D50=12.48μm, D10-D90: 6.29- 19.65μm), Ti-105-13 (Fine Powder, D50=37.31μm, D10-D90: 24.41-55.62μm), and Ti-105-27 (Coarse Powder, D50=72.15μm, D10-D90: 57.35-94.69μm), all gas atomized. The particle sizes of each powder feedstock were characterized using Microtrac S3500 particle analyzer. The particle size distribution curves for the three types of powder feedstock are shown in Fig.1. To further characterize the powders, both tap density and powder bed density were measured. For powder bed density evaluation, powder from each type of feedstock was spread by the powder spreading actions layer by layer in the system (EOS M270) with no laser melting actions.

For powder bed density evaluation, powder from each type of feedstock was spread by the powder spreading actions layer by layer in the system (EOS M270) with no laser melting actions. A total of 10mm-thick powder was spread, and all the powder from the entire build bed was subsequently2122retrieved carefully. The powder bed density was then calculated by dividing the weight of the powder by the overall build volume. A very small amount of powder was expected to be lost during the manual powder retrieval process, however due to the sufficiently large overall volume of evaluation, such error was considered negligible.a. Fineb. Mediumc. Coarse Figure 1: Powder feedstocks particle size distribution2.2.Experimental design and sample fabricationFor the thin strut geometry, diameter of 0.1mm, 0.3mm and 0.5mm were designed. The overall length of the thin struts of 0.1mm, 0.3mm and 0.5mm were set as 5mm, 8mm and 10mm, respectively.

The overall length of the thin struts of 0.1mm, 0.3mm and 0.5mm were set as 5mm, 8mm and 10mm, respectively. The main consideration of the length designs was to accommodate handling and experimental setup (e.g., sample cutting and mounting), as the length-to-diameter ratios of all the designs were considered sufficiently large (>20) for the struts to be considered slender beams. All the thin struts were orientated at 75° (in relation to x-y plane), which was arbitrarily selected. Fig. 2 shows the design and as-built thin strut samples. For the thin strut samples, encircling wall features were designed in order to provide protection to the rather fragile sample during the post- fabrication cutoff. The samples were fabricated by the EOS M270 system. Multiple levels of laser power (90W, 100W, 110W) and scan speed (360mm/s, 400mm/s, 440mm/s) were investigated, as shown in Table 1.

The samples were fabricated by the EOS M270 system. Multiple levels of laser power (90W, 100W, 110W) and scan speed (360mm/s, 400mm/s, 440mm/s) were investigated, as shown in Table 1. The choice of the baseline parameter level (80W +400 mm/s) was based on the default support scan strategy for Ti6Al4V with the EOS M270 system. Other process parameters were kept constant but are also noted here due to their relevance in the fabrication process. The layer thickness and hatch spacing distance of 30µm and 100µm were employed, which are both default for the processing of Ti6Al4V in the system. Both contour2123scan and hatch scan were enabled for all the samples. The hatch spacing scan of x-direction with 45-degree rotation on exposure was set. The process experimental design is shown in Table 1. For each setting, 5 replicates of the thin strut sample were fabricated.a. Isometric Viewb. Top Viewc. Sectional side viewd.

The process experimental design is shown in Table 1. For each setting, 5 replicates of the thin strut sample were fabricated.a. Isometric Viewb. Top Viewc. Sectional side viewd. Sample after farbicationFigure 2: (a-c) Thin Strut CAD designs and (d) actual sample after fabricationTable 1: Experimental design for process parameter settingsPowder Feedstock Thin Strut Diameter Powder Layer Thickness Hatch Scan Spacing Part OrientationNotationLaser Power (W)Fine, Medium, and Coarse 0.1 mm, 0.3 mm, and 0.5 mm 30 µm 100 µm 75° Volume Energy Density (J.mm-3)Scan Speed (mm/s)P1 (Baseline) P2 P3 P4 P5100100 100 90 110400360 440 400 40083.3392.59 75.76 75.00 91.672.3.Sample preparation and characterizationAfter fabrication, the samples were carefully cut from their support structures using an Allied TechCut 5 slow cutter. To prevent damage or loss, the chambers of the thin strut samples were filled with candle wax prior to cutting, as shown in Fig. 3.

To prevent damage or loss, the chambers of the thin strut samples were filled with candle wax prior to cutting, as shown in Fig. 3. The wax was removed afterwards by simply heating the samples.(a) liquid wax-filled sample(b) after solidification of waxFigure 3: Thin Strut Sample retrieval using wax2124The as-built samples were subjected to further characterization without stress-relieving. Thickness measurement was carried out using digital optical microscope (Olympus MX51) by taking three locations along the length of the samples and measuring the thickness from the two external boundary points within focus. One of the replicates of each sample design was grinded along the x/y-z plane and mounted with epoxy resin into metallurgical samples for porosity and microstructural analysis.

One of the replicates of each sample design was grinded along the x/y-z plane and mounted with epoxy resin into metallurgical samples for porosity and microstructural analysis. For the metallurgical sample preparation, the samples were carefully grinded to roughly half-planes by a Pace Technologies NANO 2000T sample grinding/polishing machine using 180/240 and by a semi-automated Allied MetPrep 3 Grinder/Polisher using 400- grit sandpapers. After achieving the desired sample planes, the samples were further polished with a polishing grit number sequence of 600, 800 and 1200 grit sandpapers, followed by 6µm and 1µm diamond suspension polishing with polishing lubricant in the same Grinder/Polisher at 100-150 rpm. Polished samples were first subjected to porosity characterization, including areal porosity and areal pore size measurements. For each sample, the porosity and pore sizes were measured at three different build heights (near support, middle, and large height).

For each sample, the porosity and pore sizes were measured at three different build heights (near support, middle, and large height). Image analysis-based porosity characterization was carried out using ImageJ software. While the gas pores and keyhole pores are typically spherical in shape, lack-of-fusion flaws can be large and irregular in shape [29]. A study performed by Gong et al. on defect morphology in Ti-6Al-4V parts manufactured using L-PBF process showed the defect circularity ranging from approximately 0.10 – 1.0 for process settings at over-melting zone [30]. With the aim to capture defects as small as single-micron size and the ones with low circularity, the pore size of greater than 2µm2 and circularity of 0.10-1.0 were used as identification thresholds. The lower limit of defect cross-sectional area of 2µm2 corresponds pores with ~ 1.4µm diameter with a circular cross-section which was identified to be reasonable cut-off limit based on the experimental PSD results.

Afterwards, the samples were then etched with Kroll’s reagent for microstructural analysis using cotton swab rubbing. The exact etching times varied from approximately 25 to 35 seconds depending on the reagent available on the sample during swab rubbing and sample size. The completion of etching was determined by frequently inspecting the microstructural image quality under the optical microscope. The etched samples were subjected to grain width measurement using linear interception following ASTM E112-13 (Heyn Lineal Intercept Procedure) [31] using the digital optical microscope. It is also noted that for the thin strut samples the physical exterior boundaries themselves also count as grain boundaries, which differs slightly from typical procedures.

It is also noted that for the thin strut samples the physical exterior boundaries themselves also count as grain boundaries, which differs slightly from typical procedures. Since the large-size grains mostly exhibit columnar morphology with orientations generally along the z-direction (build direction), the intercept lines were selected to be approximately perpendicular to the z-direction.The remaining 0.3 mm and 0.5 mm samples were subjected to a three-point bending test to evaluate their mechanical properties, due to the difficulty of sample gripping for typical tensile testing. The 3-point bending test setup is shown in the Fig. 4. The mechanical testing was performed on the Mark-10 ESM306 tensile testing machine with 500N load cell and a custom 3- point bending fixture, under a constant loading rate of 0.50 mm/min and span length of 4mm. The sample was slightly preloaded to ensure proper contact between the upper anvil and test sample.

The sample was slightly preloaded to ensure proper contact between the upper anvil and test sample. The equipment lowest loading rate was used such that the kinematic effects such as slippage can be minimized. The samples were tested till the failure which was identified by 70% drop in the applied load.2125(a)(b)Figure 4: Three point being test setup (a) before testing (b) after testing3. Results and discussion 3.1.Feedstock powder characteristicsThe densities of the three types of feedstocks are listed in Table 2. Considering the Ti6Al4V material density to be 4.43 g/cm3, tap density and powder bed density were normalized. The fine powder exhibits the highest tap density but the lowest powder bed density. This is likely due to the enhanced inter-particle interactions at reduced particle sizes, which causes low flowability but could be overcome with additional powder fluidization mechanisms such as vibration.

The relatively broad relative PSD range of the fine powder might also contribute to the high tap density of the powder feedstock. Although the absolute particle size range for the fine powder is narrow (~14μm between D10 and D90), the percentage size range is rather significant (D90 approximately 300% of D10). Consequently, the void-filling effect from polydisperse powder packing is more pronounced [32].

Consequently, the void-filling effect from polydisperse powder packing is more pronounced [32]. On the other hand, the medium feedstock exhibits the lowest tap density and highest powder bed density, which might be partially explained by its good flowability and spreadability, as well as the relatively narrow relative PSD range.Table 2: Different Powder feedstock densitiesPowder FeedstockFineTap density (g/cm3) 3.29Normalized Tap density0.743Powder Bed density (g/cm3) 2.33Normalized Powder Bed density0.526Medium2.860.6462.790.630Coarse3.230.7292.660.6003.2.Fabrication success rateThe success rates of the thin strut fabrication under each design condition were estimated by the percentage of successfully fabricated samples. Due to limited sample replication number (5), the success rates were estimated in rather crude scales.

Due to limited sample replication number (5), the success rates were estimated in rather crude scales. However, it was expected that such information could still provide some insights into the effects of different design input variables. As shown in Fig. 5a, the success rate for 0.1mm thin struts are higher with fine powder feedstock2126whereas for 0.3 mm and 0.5 mm thin struts coarse powder feedstock showed highest success rate. The success rate of the 0.1mm strut is higher with the fine powder feedstock, which might be resulted from the reduced variability of melt pool-powder interaction due to the larger average numbers of powder particles involved. On the other hand, for 0.3mm and 0.5mm struts, the lower powder bed density of the fine powder feedstock likely become a significant factor for defect generation and consequently lower success rates.

On the other hand, for 0.3mm and 0.5mm struts, the lower powder bed density of the fine powder feedstock likely become a significant factor for defect generation and consequently lower success rates. It could be reasonably speculated that for the thin feature fabrication, refined powder could favor the stability of melt pool evolution, which is more critical for features with shorter melt pool continuity (e.g., thin struts of small dimensions). Although the powder feedstocks investigated in this study all exhibit single-modal PSD, it could be expected that a feedstock with bi-modal PSD might be optimal for the fabrication of thin struts, as the fine powder serves to stabilize the melt pool evolution while the overall flowability/spreadability of the feedstock is ensured.Fig.5b–d provides further insights on the effect of energy density, scan speed and laser power on the success rates of thin struts. The thin struts exhibit higher success rate at lower energy density level.

The thin struts exhibit higher success rate at lower energy density level. A study conducted by Gong et al. [33] on bulk features identified the energy density window of approximately 35.34-60.78 J/mm3 for fully dense parts. Promoppatum et al. [34] recommended a similar energy density window of 36.9-87 J/mm3 for fully dense part. The energy density levels investigated in this study were set near the high end of the suggested range. In addition, due to the kinetic inertia effect with the mechanical laser deflection system (i.e., the galvanometer system), the actual dwell times of laser within the processed regions is likely higher than the nominal values, which corresponds to higher-than-nominal input energy density. This might partially explain the observed trend in this study.

This might partially explain the observed trend in this study. Except for 0.1mm samples, decrease in laser power and increase in scan speed improved the fabrication success rate (Fig.5c– d) as they result in lower energy densities.2127(a)(b)(c)(d)Figure 5: Success rate of thin struts (a) all samples (b) at different conventional energy densities (c) at constant laser power of 100W (d) at constant scan speed of 400 mm/s. The red ‘X’ represents complete missing sample.3.3.Effects on dimensional accuracyFig. 6 shows the geometrical accuracy of thin struts feature in terms of absolute error and relative error (%). Absolute error is the different between the measured strut diameter and designed strut diameter, whereas relative error is the percentage of absolute error with respect to the designed strut diameter. A trend of increase in dimensional error (i.e., decrease in dimensional accuracy) with increase in feedstock size can be observed.

A trend of increase in dimensional error (i.e., decrease in dimensional accuracy) with increase in feedstock size can be observed. Although the energy density has low significance, indication of higher dimensional accuracy was identified at lower energy density. Higher dimensional accuracy was observed with the increase in dimension. The absolute errors for fine and coarse feedstock samples appear to converge to approximately 90µm and 150µm respectively. With the utilization of optimum process settings, it is believed that the fine powder samples can generate higher success rate and dimensional accuracy with thin features. The fuse of partially sintered un-melted powder to the melted powder on the part surface due to insufficient heat input on the edge borders of samples is expected to be the main cause for the dimensional error, and exerting higher impact on lower dimension samples as such effect is expected to be similar to samples regardless of their dimensions [35-38].

The present study utilized high input energy density that leads to over-melting, thus it will be potentially beneficial to investigate lower energy density levels in future studies. As current study energy2128densities are in over-melting zone, the decrease in energy density in generally favorable for reduced dimensional error. In addition, coarse powder feedstock with largest powder particle size appears have lowest geometrical accuracy, which implies significant role of powder feedstock type geometrical accuracy of thin features as compared to investigated range of energy density in. Table 3 lists the optical microscope images of thin struts fabricated with different powder feedstocks.

Porosity from processing is mainly due to lack-of-fusion or keyholing [40, 41]. Keyholes pores are formed at higher energy densities where the vapor bubbles are trapped within the melt pool, whereas lack of fusion pores are formed at lower energy density where some regions remain unmelted [40]. Due to the higher -than-nominal energy density setting parameters in this study, the majority of pores were expected to be of keyhole type with circularity close to 1, which was verified as shown in the Fig.7. Based on the single-track studies performed using Ti6Al4V powder, the keyhole pores diameter ranges from 8µm- 127µm (i.e., areal pore size range: 50µm2 – 50,671 µm2) [40, 42, 43], and most of the pores identified in this study areal pore size are within this range (Fig.8b). Furthermore, the Fig.7d-f clearly shows the increase in number of defects with increase in input volumetric energy density.

Furthermore, the Fig.7d-f clearly shows the increase in number of defects with increase in input volumetric energy density. This increase in defects is expected to be cause for lower success rate of thin struts at higher energy density levels. The highest areal porosity and pore sizes was observed at the input energy density of 92.13 J.mm-3. The statistical analysis performed on the available data points showed that powder feedstock, strut dimension and volumetric energy density have significant (p < 0.05) effects on areal porosity, whereas only strut dimension and volumetric energy density have significant (p < 0.05) effect on pore size. Both average areal porosity and pore size were found to be consistent (p > 0.05) along the build direction.

Both average areal porosity and pore size were found to be consistent (p > 0.05) along the build direction. Fig.8a-b show the average areal porosity and average areal pore size of thin strut samples.2130Figure 7: (a-c) Areal Porosity of thin strut samples at baseline process settings (a) 0.1mm fine powder feedstock, (b) 0.3 mm medium powder feedstock, and (c) 0.3 mm coarse powder feedstock. (d-f) 0.5 mm medium powder feedstock at different energy density (d) 75.38 J.mm-3 (e) 83.33 J.mm-3 (f) 92.13 J.mm-3(a)(b)Figure 8: (a) Average areal porosity and (b) Average pore size of thin strut features at different energy density. Due to low success rate, no Tukey pairwise comparison was performed. Graphs does not include the missing samples.2131Zhang et al. [35] and Delcuse et al. [44] reported decreasing trend of porosity with decrease of thin strut dimensions.

Graphs does not include the missing samples.2131Zhang et al. [35] and Delcuse et al. [44] reported decreasing trend of porosity with decrease of thin strut dimensions. For the feature dimension range in this study, similar decreasing porosity trend can be observed with decrease in strut diameters. Average pore size also showed decreasing trend with decrease in strut diameters. Overall, the pores size appears to exhibit broad ranges for individual samples whereas the effect of the thin feature dimension is relatively insignificant. Further checking with the scan paths generated by the EOS PSW3.5 control software showed that only contour scan was active in the 0.1 mm thin strut samples, whereas both contour and hatch scan were active in the 0.3 mm and 0.5mm samples. As the 0.1 mm samples are not exposed with hatch scan, the total areal input energy density is expected to be lower than for the 0.3 mm and 0.5 mm samples.

As the 0.1 mm samples are not exposed with hatch scan, the total areal input energy density is expected to be lower than for the 0.3 mm and 0.5 mm samples. On the other hand, the hatch scan for 0.3 mm and 0.5 mm samples could potentially leads to excessive heat input due to the frequent laser acceleration/deceleration and turning [41]. The input energy in this study is in higher than nominal value and excessive heat input at the turnaround locations could further exacerbate quality issues such as dimension, porosity, and pore size for the 0.3mm and 0.5 mm strut samples.3.5.Effects on microstructure and grain widthThe grain microstructure of thin features fabricated using different powder feedstock are shown in the Fig. 9a. Columnar grains growing along the build direction were observed. Similar to the study by Zhang et al. [45], the grain width in the strut center along the diameter direction were identified to be larger than those located near the strut boundaries (Fig.9b).

Similar to the study by Zhang et al. [45], the grain width in the strut center along the diameter direction were identified to be larger than those located near the strut boundaries (Fig.9b). Small grains at the feature surface nucleate from adjacent powder particles at the part surface in the power bed and grow towards the center following the maximum temperature gradient [12]. Since the grain growth at surface is inhibited by the higher cooling rate at surface, the grain width at the strut boundaries were smaller than in the interior regions. The graph in the Fig. 9c indicates the variation of grain width of thin struts. For current study, the results clearly indicate the higher influence of thin strut dimension as compared to powder feedstock type and input energy density. The cooling effect of loose powders at strut boundaries are similar irrespective of the strut dimension.

The cooling effect of loose powders at strut boundaries are similar irrespective of the strut dimension. However, with increase in strut diameter, the strut boundary cooling effect reduces, promoting the formation of more number of grains in the center. However, the grain width of the center grains is larger than boundary grains. Thus, higher number of larger center grains in larger diameter thin strut results in overall grain width increment. More regular grain morphology towards the center interior of the feature is also partially attributed to the continuous growth of columnar grain due to heterogeneous grain growth. Similar to porosity, for 0.3mm samples the increase in particle size distribution resulted in decrease in grain width. On the other hand, the grain width of 0.5mm were largely consistent across different powder feedstocks.

On the other hand, the grain width of 0.5mm were largely consistent across different powder feedstocks. This might imply that as the feature dimension increases, the solidification and cooling process within the L-PBF-AM gradually achieves stabilized status with the bulk of the features, and therefore the dependency of grain width on the powder feedstock size gradually diminishes.2132(a)(b)(c)Figure 9: (a) Thin Strut Grain microstructure of different powder feedstock (b) 0.5mm medium powder feedstock thin strut sample showing interior grain boundaries indicated with red dots (c) Average grain width of thin strut samples at different energy density3.6.Effects on feature mechanical strengthThree-point bending test were employed to evaluate the flexural modulus, flexural yield strength and ultimate flexural strength of different samples, and the results are presented in Fig.10. Sun et al.

Sun et al. [46] studied the effects of build direction on mechanical performance of L-PBF additively manufactured Ti6Al4V at input energy density level of 58.33 J.mm-3, and reported2133the ultimate flexural strength to be between 948 MPa (at 0° build orientation) to 1026 MPa (at 90° build orientation). In this present study, 0.5 mm thin strut fabricated with coarse powder feedstock showed lowest value of average mechanical properties (i.e., flexural modulus of 13.54 GPa, flexural yield strength of 737.34 MPa, and ultimate flexural strength of 1392.07 MPa) at input energy density level of 75.38 J.mm-3. The yield point was identified using the derivative approach developed by Christensen [47]. The stress-strain curve within the approximate elastic- plastic region was first fitted with 5-order or 6-order polynomial function, and the yield point was then identified as the first local maxima point of the 2nd derivative of the stress-strain function.

The overall yield strength calculation approach is briefly illustrated in the Fig.11. Based on the calculated individual sample yield strength, the flexural modulus was then calculated from the identified elastic regions.(a)(b)(c)Figure 10: Mechanical properties of thin strut samples at different energy density (a) Flexural modulus (b) Flexural Yield Strength and (c) Ultimate Flexural Strength2134(a)(b)(c)Figure 11: Yield Strength analysis from stress-strain curve based on derivative approach (a) Stress-strain curve (b) First derivative of stress-strain curve Vs strain (c) Second derivative of stress-strain curve Vs strainFrom the results, the effect of the strut diameter was more significant on flexural modulus and flexural yield strength, and less significant on ultimate flexural strength.

Comparing between the medium and coarse powder feedstock, the coarse powder feedstock appears to negatively affect the mechanical properties of the struts, which is particularly significant with larger-size struts (0.5mm). Such observation was in good agreement with the observed trends with porosity and pore size discussed earlier, which increase with increase in strut size and powder particle size distribution. On the other hand, despite the expectation of lower mechanical properties at highest energy density level (i.e., 92.13 J.mm-3), in this study the effect of energy density for a given strut diameter were observed to be less significant compared to the powder feedstock type and strut dimension.21354.

ConclusionsThe thin strut features of dimensions varying from 0.1mm to 0.5mm were fabricated using L- PBF-AM process and the effect of different Ti6Al4V powder size (fine, medium, and coarse) at three different volumetric energy density i.e., 75.38 J.mm-3, 83.33 J.mm-3 and 92.13 J.mm-3 were studied. A systematic sequential methodology was developed and implemented for careful characterization and testing of the thin struts. Since the present study utilized higher-than- nominal input energy density, utilization of lower input energy density close to nominal values are recommended for higher fabrication success rate. Within the range of process parameters studied in this work, bimodal powder size distribution is expected to yield better fabrication success rate. More importantly, both thin strut dimension and powder feedstock size distribution were highly influential factor from strut quality aspect.

More importantly, both thin strut dimension and powder feedstock size distribution were highly influential factor from strut quality aspect. Based on the results on fabrication success rate, dimensional accuracy, porosity, and pore size, microstructure grain size and different mechanical properties, coupling effect with compound PPG-M relationships were observed for thin strut features. Therefore, this study elucidates the necessity of further in-depth studies on small dimensional scales for L-PBF-AM process such that the knowledge base on the thin features L-PBF-AM process can be re-established.Declaration of Competing InterestThe authors declare no conflict of interest.AcknowledgementsThis project was partially supported by the Praxair Surface Technologies TruForm Ambition Grant.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceThe Effects of Powder Recycling on the Mechanical Properties of Additively Manufactured 17-4 PH Stainless SteelPooriya Dastranjy Nezhadfar1, 2, Arash Soltani-Tehrani1, 2, Amanda Sterling1, 2, Nicholas Tsolas1,2, Nima Shamsaei1, 2,*1National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849, USA2Department of Mechanical Engineering, Auburn University, Auburn, AL 36849, USACorresponding Author: shamsaei@auburn.eduAbstractThe booming interest in Additive Manufacturing (AM), is seeing a rising number of industries and research entities adopting this technology into their manufacturing practices. Of particular interest is Laser Powder Bed Fusion (L-PBF) process, a common AM method for fabricating metallic components. However, one obstacle is the high cost of powder feedstock.

Of particular interest is Laser Powder Bed Fusion (L-PBF) process, a common AM method for fabricating metallic components. However, one obstacle is the high cost of powder feedstock. A popular tactic to offset this cost is to reuse the powder between prints, but there is no in-depth understanding of how the powder feedstock may change or affect the mechanical properties of the produced parts. By incorporating unique powder/part characterization methods, this study quantifies the rheological properties of continually recycled 17-4 precipitation hardening (PH) stainless steel (SS) powder through successive printing of mechanical test specimens.

The AM specimens are subjected to tensile tests, to correlate mechanical behavior to changing powder quality, including particle size/shape distribution, flowability, and density.Keywords: 17-4 PH stainless steel; Laser powder bed fusion; Powder characterization; Powder recycling; Mechanical properties; Tensile behaviorIntroductionAdditive manufacturing (AM) has become a well-known layer-based manufacturing technique which helped in overcoming several limitations inherent to subtractive manufacturing methods – allowing end users the ability to fabricate near-net shaped parts directly from computer aided design (CAD) models [1, 2]. AM facilitate the production of parts with complex geometries which are unobtainable through the traditional manufacturing methods. Laser powder bed fusion (L-PBF) is one of the most common AM processes for fabrication of metallic components [3].

Laser powder bed fusion (L-PBF) is one of the most common AM processes for fabrication of metallic components [3]. In this process, parts are fabricated by repeatedly melting and fusing a thin layer of powder by a high- powered laser source to successively build components layer-by-layer. Titanium and aluminum alloys, various grades of stainless steels and Nickel-based super alloys are the most implemented materials by this method [4]. L-PBF systems are preferred due to their ability to produce finer resolution and higher quality parts compared to other AM methods [5].Although AM has its advantages, there are some issues, which must be thoroughly investigated before these parts are being used as critical load bearing parts. Various defects1292inherited from the process such as porosity, lack of fusion voids and surface roughness are the main drawbacks, which may affect the mechanical performance of the parts, especially under cyclic loading.

Utilizing appropriate process parameters, build orientation and powder conditions have been reported as important parameters to control or reduce the aforementioned issues [6]. Considering the powder is one of the most crucial elements in the L-PBF method, many studies in the past have examined how powder characteristics effect the mechanical behavior of fabricated parts. Recently, morphological characteristics comprising size, shape and surface roughness have been studied. Irrinki et al. [7] studied gas- and water-atomized powders with different mean particle sizes, and observed lower porosity and higher part densities when fabricated with finer particle sizes. Parteli et al. [8] developed a particle-based numerical simulation and reported larger surface roughness values when parts are fabricated with powders containing a larger size distribution of particles.

[8] developed a particle-based numerical simulation and reported larger surface roughness values when parts are fabricated with powders containing a larger size distribution of particles. This was attributed to the finer particles tendency to form larger agglomerates, which increases porosity as a result of larger particles inability to fill voids during packing. Sutton et al. [9] has mentioned that the flowability and layer homogeneity are highly influenced by particle size and size distribution. Moreover, the large particles and agglomerates may produce voids because of the reduction in packing density.Other studies have also investigated how recycling and reusing can change powder characteristics and its relation to the mechanical properties of fabricated parts. Powder reusing is of high importance to industries because of its economic and environmental perspective.

Powder reusing is of high importance to industries because of its economic and environmental perspective. Currently, using either the un-melted metal powder from the previous builds or mixing with virgin powder are potential strategies. However, it has been indicated that recycling powders comprised of powder particles that have been exposed to the building chamber conditions, with condensate powder particles splattered from the melt pool can have an effect on the fabrication fusion process, which can lead to variation in the overall material properties of the part [10].Debate remains on the methodology and effects of powder recycling on AM material properties. Some studies have reported no variation in mechanical properties fabricated via recycled powders [11-13]. Jelis et al. [11] studied the effects of powder recycling on the mechanical properties of AM 4340 steel fabricated through direct metal laser sintering (DMLS).

Jelis et al. [11] studied the effects of powder recycling on the mechanical properties of AM 4340 steel fabricated through direct metal laser sintering (DMLS). Virgin and once-recycled powder were used and passed through an 80 µm sieving screen in order to eliminate the overexposed and splattered condensate particles. Insignificant changes in chemical composition of powders, and tensile properties were reported. Similarly, Jacob et al. [12] reported insignificant changes in shape, particle size and particle size distribution, but an increase in the apparent density and powder bed density were observed after recycling 17-4 PH stainless steel powders up to 11 times. Although, some changes in powder rheological properties were reported, there were no significant changes in the tensile behavior of tested specimens. In addition, it has been reported that IN718 powder properties and tensile deformation behavior had no change after 14 times powder recycling [13].

In addition, it has been reported that IN718 powder properties and tensile deformation behavior had no change after 14 times powder recycling [13]. On the other hand, Tang et al. [14] reported an increase in the yield and tensile strength of Ti-6Al-4V parts after 21 times powder recycling. This was attributed to the increase in oxygen content absorbed during the fabrication process.Although many studies have investigated the powder recycling effects on the tensile behavior of fabricated parts, the fatigue performance of the parts have not been studied. As mentioned earlier, changes in powder rheological properties induce porosity and surface roughness. Many studies have reported the detrimental effects of defects such as pores, lack of1293fusion and surface roughness on the fatigue performance of AM parts [2, 3, 15, 16].

Many studies have reported the detrimental effects of defects such as pores, lack of1293fusion and surface roughness on the fatigue performance of AM parts [2, 3, 15, 16]. Therefore, understanding the powder recycling effects on the cyclic deformation behavior of AM parts is essential.17-4 precipitation hardening (PH) stainless steel (SS) has drawn significant attention recently due to its outstanding characteristics, such as high tensile/impact strength, fracture toughness and corrosion resistance at service temperatures below 300 °C. This material is widely used for structural components in aerospace, petrochemical, power plants and marine environments [17, 18]. The goal of this study is to investigate the effects of powder degradation on the microstructural evolution, monotonic and cyclic deformation behavior of 17-4 PH SS after powder recycling.

The goal of this study is to investigate the effects of powder degradation on the microstructural evolution, monotonic and cyclic deformation behavior of 17-4 PH SS after powder recycling. Moreover, identifying a measurable variable of powder properties indicative of the part quality is another goal of this ongoing study. The preliminary results of this broad powder study will be presented in the present paper. Following the introduction, the material and experimental procedures are presented. The powder characterization results and tensile deformation behavior are then discussed. Finally, important conclusions are drawn.Experimental proceduresArgon- atomized 17-4 PH SS powder with the particle size distribution (PSD) of 15-45 µm provided by LPW Technology Inc. was used for this study. The chemical composition of the virgin powder is presented in Table 1. 15 sets of prints were carried out using 80 kg of powder, with the aim of recycling the powder after each print.

The chemical composition of the virgin powder is presented in Table 1. 15 sets of prints were carried out using 80 kg of powder, with the aim of recycling the powder after each print. Figure 1 shows the full print and half print layouts. The full print layout consists of tensile, fatigue and impact test specimens, while the half print layout only includes the tensile and impact test specimens. This strategy not only minimized fabrication time, but also afforded a greater number prints (number of powder recycling times). Batch numbers 1, 5, 10 and 15 followed the full print layout, while batches in between followed the half print layout. Specimens were fabricated vertically in a nitrogen atmosphere using an EOS M290 L-PBF system. Default build parameters were used, as well as the fabrication started with unused, virgin powder (batch 1).

Default build parameters were used, as well as the fabrication started with unused, virgin powder (batch 1). The un-melted powder after each fabrication was being collected and sieved through an 80 µm screen and mixed with the original powder in the feed bin to generate the feedstock for the next batch. After mixing the powder, a sample representative of all the powder in the feed bin was obtained according to ASTM B215-15 [19] (equal increments from different depths of the powder). This process was continued after ever printed batch until the 15th print. It is worth mentioning that no virgin powder was added in between the fabrication of batches, in order to isolate the effects of powder recycling on the rheological properties and final mechanical response of the parts.To characterize the rheological properties of the powder used in fabricating batches 1, 5, 10 and 15, a Freeman Technology rheometer (FT4) machine was employed.

This instrument is capable of measuring various rheological powder properties including bulk/apparent and tapped (maximum) density, compressibility, aeration ratio, cohesion, permeability, and shear stress. This instrument quantifies the powder’s rheological properties measuring the amount of work done in moving the blade through the powder from the top to the bottom of the vessel. In the current stage of this study, only powder bulk density, aeration ratio, and cohesion were measured.1294Table 1. Chemical composition of 17-4 PH stainless steel powderCCrNiCuMnSiNbMoNO(Wt. %)0.0115.64.033.890.240.290.33<0.010.010.05(a)(b)Figure 1.

Chemical composition of 17-4 PH stainless steel powderCCrNiCuMnSiNbMoNO(Wt. %)0.0115.64.033.890.240.290.33<0.010.010.05(a)(b)Figure 1. The layout of prints: (a) full print, (b) half printSpecimens were heat treated, which included a 30-min solution treatment at 1050˚C, followed by air cooling to the room temperature, and subsequent annealing for 4-hrs at 552˚C, followed again by air cooling to room temperature.To evaluate the mechanical properties, specimens were subjected to uniaxial tensile tests at room temperature at 0.001 s-1 strain rate using MTS Landmark servo hydraulic test system with 100 kN load cells. Tests were carried out in two steps, strain-controlled and displacement controlled.

Tests were carried out in two steps, strain-controlled and displacement controlled. Since the elongation to failure of the material is higher than the resolution of the extensometer, the strain-controlled step was carried up to 0.045 strain and continued in displacement control to the final fracture by removing the extensometer. For certainty, three tensile tests were conducted for each batch to see the consistency of the results and eventually, the average is presented.Results and discussionPowder QualityFigure 2 represents the variation in density of powder samples obtained from batches 1, 5, 10 and 15. As observed, the powder density increased with the number of recycling times. This is attributed to the change in the powder particles size distribution.

As observed, the powder density increased with the number of recycling times. This is attributed to the change in the powder particles size distribution. Finer particles melted and solidified, while particles greater than 50 µm were pushed away from the build plate via the recoater arm since the pre-adjusted gap between the substrate and the recoater arm was smaller1295FeBal.than 50 µm. It has been previously reported that due to the shape, frictional forces, particle interlocking and changes in the moisture level of the powder, particles are expected to agglomerate [14]. As such, any agglomerated particles passed through the sieve screen were broken down into finer particles. This effect will narrow down the PSD with each recycling time. Any changes in PSD will affect the porosity distribution in the powder bed.

This effect will narrow down the PSD with each recycling time. Any changes in PSD will affect the porosity distribution in the powder bed. In fact, the recycled powder with the narrower PSD has a lower porosity in the powder bed leading to a higher density [9].4.34.24.144.19) lm / g ( y t i s n e D4.143.93.874.033.83.7Batch 1Batch 5Batch 10Batch 15Batch numberFigure 2. Powder density as a function of recyclingFigure 3 compares the aeration ratio of the reused powder for batches 1, 5, 10 and 15. Aeration ratio is an indication of powder sensitivity to air pockets. The aeration ratio represents how easily the gas can flow within the powders and transfer them to the building platform. As shown, the aeration ratio decreased from batch 1 to 5, and gradually increased up to batch 15.

As shown, the aeration ratio decreased from batch 1 to 5, and gradually increased up to batch 15. Accordingly, batch 5 possesses a higher sensitivity to air pockets, which represents less resistance to inter-particle forces, consequently increasing the flowability of the powder. The aeration ratio has a direct relation with cohesion, which means that powders with higher aeration ratios have higher cohesion due to frictional forces and particle interlocking. Accordingly, cohesivity of batch 5 is lower than batches 1, 10, and 15, which makes the powder fluidized and thus improves the flowability of powder over the build plate.1296140.3120.25o i t a r n o i t a r e A108640.20.150.1) a P k ( n o i s e h o C2Aeration ratio Cohesion0.05000510 Batch number1520Figure 3. Aeration ratio and cohesion of the powder over recycling timesTensile behaviorFigure 4 compares the tensile deformation behavior of specimens from batches 1, 5, 10 and 15.

Aeration ratio and cohesion of the powder over recycling timesTensile behaviorFigure 4 compares the tensile deformation behavior of specimens from batches 1, 5, 10 and 15. Figure 4a represents the strain-controlled step up to 0.045 strain before the extensometer was removed. Figure 4b shows the stress-displacement curve up to the failure. In addition, the tensile behavior of the different batches is compared in detail in Table 2. At first glance, it is observed that the tensile deformation behavior has not been affected significantly by incremental change in powder recycling times. However, details in Table 2 reveal a 20% increase in reduction of area ) for 5th batch compared to 1st (RA) and approximately 23% increase in true strain at fracture ( batch. On the other hand, it seems that incremental changes in powder recycling after the 5th batch has led to a slight decrease (~ 6-8%) in RA and.

On the other hand, it seems that incremental changes in powder recycling after the 5th batch has led to a slight decrease (~ 6-8%) in RA and . In other words, the ductility has increased after 4 iterations in powder recycling (5th batch) and decreases after a higher numbers of iteration times 𝜀𝜀𝑓𝑓 (10th and 15th batch). The increase in ductility of 5th batch is attributed to its lower cohesion compare to other batches (Figure 3). As mentioned earlier, lower cohesion means the powder possess higher flowability, which would affect the contiguity of the part. Accordingly, increase in the cohesion of powder would decrease the ductility of the part.

Accordingly, increase in the cohesion of powder would decrease the ductility of the part. Moreover, a slight decrease in strength is observed for the 15th batch.𝜀𝜀𝑓𝑓129714001400(a)(b)1200120010001000) a P M800) a P M800( s s e r t S600( s s e r t S600400batch 1400batch 1batch 5batch 5200batch 10200batch 10batch 15batch 150000.010.02Strain0.030.040.0500.511.5 Displacement (mm)22.5Figure 4. Tensile behavior of batches 1, 5, 10 and 15: (a) stress-strain curve up to the 0.045 strain, (b) stress-displacement curve up to the failureTable 2. The details of tensile test results of batches 1, 5, 10 and 15 5th 35.40 0.431st 29.50 0.3510th 34.70 0.42Print number RA%15th 33.00 0.40𝜺𝜺𝒇𝒇ConclusionsThis study was carried out to investigate the changes in rheological properties of recycled powder after 15 sets of prints and the subsequent effects on the mechanical properties of the parts.

The goal of this study was to propose a specific change in powder rheological properties which causes variation in mechanical properties of the parts. To this end, 15 batches of specimens were fabricated continuously with a certain amount of powder without adding unused virgin powder. The remaining powder after fabrication of each batch was sieved and used for the next subsequent batches. Powder rheological properties were analyzed for print numbers 1, 5, 10, and 15, and compared to the unused powder. Moreover, the mechanical properties of the specimens belonging to the above mentioned set of prints were studied based on tensile mode of deformation. The following conclusions can be made based upon the results:1. The apparent/bulk density of the powder increased with recycling times. This is attributed to the change in the PSD after recycling-fabrication-recycling loop. PSD becomes narrower after each recycling time.2.

This is attributed to the change in the PSD after recycling-fabrication-recycling loop. PSD becomes narrower after each recycling time.2. Powder cohesion decreased from batch 1 to 5 and then increased from batch 5 to 10 and plateaued at 15.3. Powder flowability of batch 5 was the highest amongst all batches examined. 4. Powder recycling up to 5th batch resulted in a 20% increase in RA and 23% increase in . This is attributed to the decrease in cohesion and an increase in the flowability of powder up to this level of recycling, which also affects the structural integrity of the part.𝜀𝜀𝑓𝑓12983AcknowledgementsThis work is sponsored by the Naval Air Systems Command (NAVAIR)References[1] T.D. Ngo, A. Kashani, G. Imbalzano, K.T.Q. Nguyen, D. Hui, Additive manufacturing (3D printing): A review of materials, methods, applications and challenges, Composites Part B: Engineering 143 (2018) 172-196. [2] A. Yadollahi, N.

As the powder is spread by the recoater, the packing state may not be uniform on the powder bed, giving rise to location-dependent part performance. In addition, as the powder is reused (a common practice in AM industry), its characteristics continuously evolve, causing the defect content to change from build to build. This study aims to investigate the effects of powder reuse and part location on powder characteristics as well as the defect structure of the parts. Results indicate powder reuse in an L-PBF system may reduce the number of defects in the as-fabricated parts due to the superior packing state of reused powder.

Results indicate powder reuse in an L-PBF system may reduce the number of defects in the as-fabricated parts due to the superior packing state of reused powder. Part density was also found to be location-dependent, with more defects near the gas outlet.Keywords: Additive Manufacturing, L-PBF, Powder Recycling, Ti-6Al-4V, Particle Size Distribution (PSD), Location Dependency, Packing State, Powder FlowabilityIntroductionMetallic powder is used as the material feedstock in laser powder bed fusion additive manufacturing (L-PBF AM). Therefore, different powder properties can influence the performance of additively manufactured (AMed) parts. For instance, the flowability of non-spherical particles can be inferior due to high inter-particle friction and mechanical interlocking [1–4], and vice versa. As a result, more flowable particles tend to be deposited farther away from the feed bin and the less flowable ones tend to be deposited closer.

As a result, more flowable particles tend to be deposited farther away from the feed bin and the less flowable ones tend to be deposited closer. As a result, a greater location dependency of defect content and mechanical properties are expected for powders with more non-spherical particles [5].Another powder characteristic that can affect its physical behavior is the particle size distribution (PSD). Fine particles are well known to result in high theoretical powder bed density (PBD) and coarser ones can give rise to higher flowability [6–9]. The reason is that finer particles have typically larger surface-to-volume ratios resulting in higher friction and interlocking [10]. On the other hand, fine particles can fill the voids between coarse particles to give a denser particle arrangement [11].Powder chemical composition is another critical factor that can affect the final part performance.

For instance, the powder can agglomerate with moisture resulting in inferior1093flowability and packing state [11], thus higher defect content in the resulting parts. In this condition, inferior mechanical performance such as fatigue resistance and ductility can be expected. During fabrication, the concentration of interstitial elements such as oxygen and nitrogen can also change. The increase of oxygen in Ti-6Al-4V (Ti64) powder has been reported to result in lower ductility and higher material strength [12].Therefore, powder reuse which can considerably influence the powder characteristics including PSD and particle morphology due to the recycling practice (i.e., top-up or continuously reusing) can be expected to alter the mechanical performance of the AMed parts [13]. In this study, the powder feedstock was continuously reused and the microstructure and defect content of AMed parts were investigated.

In this study, the powder feedstock was continuously reused and the microstructure and defect content of AMed parts were investigated. In addition, parts were placed in four different quadrants to capture any possible location dependency of microstructure and mechanical performance. In the next sections, the experimental program will be provided which will be followed by the experimental results and discussion. Finally, some conclusions will be drawn.Experimental ProgramPlasma-atomized Ti64 Grade 5 powder from AP&C, a GE Additive Company with an initial PSD of 15-53 µm was used as feedstock. After each fabrication, the powder was sieved using an 80-µm mesh and reused for the next iterations. The powder was reused 8 consecutive times without adding any virgin (i.e., unused) powder. It needs to be mentioned that each fabrication was 100 hours giving a total of 800 hours of powder exposure to the fabrication environment. To the authors’ knowledge and based on Ref.

It needs to be mentioned that each fabrication was 100 hours giving a total of 800 hours of powder exposure to the fabrication environment. To the authors’ knowledge and based on Ref. [13], this is one of the research studies with longest hours of fabrication on the reuse of Ti64 in L-PBF. The same layout was used for all fabrications (see Fig. 1) which consisted of 13×13×97 and 9×9×83 mm3 square blocks to be later machined into fatigue and tensile specimen geometries according to ASTM E466 and E8 [14,15]. The blocks were distributed into four major quadrants, with each quadrant containing some high- strain fracture specimens intended for another study.The build plate also consisted of some X-ray computed tomography (XCT) coupons (see Fig. 2(a) for the geometry) which were designed to capture the effects of powder reuse and location on the resulting defect content.

2(a) for the geometry) which were designed to capture the effects of powder reuse and location on the resulting defect content. These coupons were placed in the northeast (NE), north (N), northwest (NW), east (E), center (C), west (W), southeast (SE), south (S), and southwest (SW). two powder containers for PBD measurement (see Fig. 2(b) for the geometry) at different locations were placed in the front (i.e., near to the powder feed bin) and back (i.e., away from the feed bin). Lastly, one microstructure specimen per quadrant was placed on the build plate for further microstructural analyses.Parts were manufactured using an EOS M290, an L-PBF machine using Ti64 Grade 5 process parameters (Ti64Grade5_040_HiPer). The major infill process parameters were laser power (P) of 280 W, scanning speed (V) of 1300 mm/s, hatching distance (h) of 120 µm, and 40 µm layer thickness (t) resulting in an energy density (E) of 44.87 J/mm3 (E=P/Vht).

Moreover, a laminar flow nozzle was used for fabrication as recommended by the machine manufacturer for processing Ti64 Grade 5. Before each fabrication, the powder was sampled based on ASTM B215 from the feedstock, front, and back locations [16].In addition, some heat-affected powder was collected from the spatter-rich regions (i.e., near the inert gas outlet as well as the inside and on top of the gas outlet nozzle), as specified in1094Ref [17], to track the changes in spatter particles, including their size and population, with continuously reusing the powder. It needs to be specified that all parts were fabricated on 3 mm support structures for convenient removal from the build plate. After fabrication, the XCT and microstructure coupons, as well as the powder containers were removed from the build plate in non-heat-treated (NHT) condition and the remainders of the parts were stress-relieved at 704 °C for 1 hour followed by furnace cooling in an argon environment.

The 704 °C SR temperature was selected based on the previous studies on Ti64 [6,7,18,19].Fig. 1 (a) The top view of the build layout used for fabrication, and (b) parts on the build plate after fabrication.While the eight times of powder reuse was not significant compared to reuse practices of other practitioners the total time that the powder was exposed to the chamber environment was substantial (100 hours for each built totaling 800 hours) [13]. The XCT coupons were scanned from print 1 (i.e., virgin powder), print 4 (i.e., 3x reused powder), and print 8 (i.e., 7x reused powder) via Zeiss Xradia 620 Versa XCT system to include the first, middle, and last fabrication sets. The imaging conditions were: 4x objective lens, 140 kV voltage, 150 µA current, voxel size 5 µm. To avoid false defect detection, any object detected via intensity thresholding smaller than 10 µm was discarded as defects smaller than 10 µm have usually less probability to result in fatigue failures [20].

The image slices obtained from XCT were then reconstructed using the Zeiss built- in software, and the defects were evaluated using ImageJ.1095(a)(b)Fig. 2 The geometry of (a) XCT coupons used for defect characterization and (b) powder containers for PBD measurements.The PBD density was also evaluated with a high-precision scale to capture the weight of the stored powder in the capsules. For powder characterizations, an FT4 powder rheometer was used to obtain powder rheological properties. The analyses included shear stress tests according to ASTM [21] bulk and tapped density tests, and compressibility. Lastly, the morphology and surface chemistry of powder particles were obtained using a Zeiss Crossbeam 550 scanning electron microscope (SEM) and the onboard Oxford energy-dispersive X-ray spectroscopy (EDS) detector. Finally, the hardness for the half-built specimens was measured using a Rockwell C-scale indenter (at 150 kg) with LECO LCR 500.

Finally, the hardness for the half-built specimens was measured using a Rockwell C-scale indenter (at 150 kg) with LECO LCR 500. The tests were repeated at least 3 times per condition.Results and DiscussionThe scanning electron microscopy images of particle morphology were obtained for the virgin, 3x-, and 7x- reused powder batches and the associated spatter-rich powder are presented in Fig. 3. The powder particles were highly spherical even after reusing the powder 7 times which was consistent with the observations reported in Ref. [22]. No considerable difference in morphology could be seen.

[22]. No considerable difference in morphology could be seen. A limited number of irregularly shaped and elongated particles were detected in all the powder batches (i.e., virgin, 3x- and 7x-reused) which can be due to the fabrication or powder handling.Some particles in the reused batches showed re-solidification marks on the surface (i.e., shown by yellow arrows in the figure) which can be due to melting and re-solidification inside the chamber environment as explained in Ref [23]. Such particles were abundant in the spatter-rich powder batches and generally in the form of very large agglomerates (i.e., clusters of particles) which is typically due to the particles colliding with each other and forming agglomerates while in flight in the chamber environment [24]. The fact that only a limited number of such particles were detected in the reused batches indicates the effectiveness of the sieving process.

The fact that only a limited number of such particles were detected in the reused batches indicates the effectiveness of the sieving process. It needs to1096be noted that the PSD characterization was not conducted and the observations on morphological aspects were solely based on SEM images.Fig. 3 Morphology of the Ti64 powder particles in virgin, 3x-reused, and 7x-reused conditions as well as the spatter powder after fabrications collected from spatter-rich regions. The scale bars in all images are 50 µm. The yellow arrows indicate the particles with remelting lines and yellow dashed circles enclose agglomerates. Red arrows also are pointed toward the smaller particles that are not affected by the fabrication process.The majority of spatter-rich powder particles were spherical as the spatters have enough time to re-solidify in the chamber environment under argon atmosphere which can result in a very large powder size.

The remaining more spherical particles with smaller sizes may be the ones unaffected by the fabrication process. Some of these particles are not removed from the chamber and can land near the spatter-rich regions in the EOS M290.Indeed, the sieving process can remarkably contribute to the powder quality by removing such large agglomerates. These agglomerates are also reported to consist of oxide layers that can decompose during fabrication and increase the oxygen content of the parts [23,25]. In addition, the irregular shape of these particles along with the oxide film around them can decrease the heat absorptivity, giving rise to partially sintering to the side surfaces of the specimens and increase the surface roughness [17].

Lastly, some of these particles, if not properly removed, can deposit on the top surface of the specimens yielding to balling defects and disturbing the uniform powder layer distribution and more defects in the as-fabricated parts [20].The second powder characterization step was to investigate the changes in rheological properties due to powder reuse. Powder compressibility, cohesion, and aeration energy (AE) are shown in Fig. 4. Compressibility is defined as the amount of change in density when normal stress is applied. The 7x-reused batch has the lowest compressibility as shown in Fig. 4 (a), which indicates a superior packing state in the 7x-reused powder batch and is which was consistent with Ref. [4,8,26].1097Powder cohesion can represent the resistance of the powder to flow, with both being positively correlated. In L-PBF machines, the powder is distributed on the build plate via a recoater (also known as wiper or blade).

In L-PBF machines, the powder is distributed on the build plate via a recoater (also known as wiper or blade). Therefore, a powder batch with minimal resistance to shear stress and flow is desired. It is evident from Fig. 4 (b) that the cohesion decreased with reuse, suggesting a higher flowability and possibly a more uniform powder layer for the 7x-reused powder. The reasons why cohesion decreased with powder reuse are still under investigation. However, some possible factors can be the changes in PSD with powder reuse. As reported in [7], continuously reusing the powder may result in a decrease in finer particles as they are either used in the fabrication or ejected from the build plate during fabrication. In addition, large agglomerates were reported to decrease with the sieving process resulting in a narrower PSD. A narrower PSD is commonly associated with higher powder flowabilities [7,12,27].(a)(b)(c)Fig.

A narrower PSD is commonly associated with higher powder flowabilities [7,12,27].(a)(b)(c)Fig. 4 A comparison of (a) compressibility, (b) cohesion, and (c) aeration energy (AE) among the virgin, 3x- and 7x-reused powder batches.Lastly, the AE was evaluated which identifies how easily the powder can flow [28]. Therefore, powder batches with higher AE values have usually a higher tendency to agglomerate [4]. As seen in Fig. 4 (c), it seems that AE to some extent increases with powder reuse. This trend can show that there is possibly less mechanical interlocking when the virgin powder was used for1098fabrication. Therefore, fewer agglomerates should have been present in this batch. The reason(s) that why agglomeration slightly increased with powder reuse is also under investigation. A possible reason, however, can be explained by spattering during fabrication.

The reason(s) that why agglomeration slightly increased with powder reuse is also under investigation. A possible reason, however, can be explained by spattering during fabrication. As particles eject from the build plate, they have enough time to re-solidify in the chamber environment, and in some cases, collide with each other and form larger particles [23]. Although some of these large agglomerates are successfully filtered out using the sieve, some may pass through if they are smaller than 80 µm.Therefore, the reused batch with some of these agglomerates can show a higher AE or higher mechanical interlocking. A question might raise here that why the cohesion showed a decrease with powder reuse. This can be answered by the fact that in some conditions, agglomerates can move in the shape of “lumps” and show good flowabilities and low cohesion [28] whereas it can result in a nonuniform powder layer deposition.

Therefore, the PSD characterization would be indeed needed to understand why the reused powder showed a higher AE and lower cohesion.To evaluate the quality of the deposited powder layers during fabrication, PBD was assessed using the PBD capsules. The PBD evaluations, however, did not show any changes with powder reuse and they were between 2.63-2.65 g/cm3 for different powder reuse iterations. However, it was found that the PBD is slightly lower in the back location (further away from the powder feedstock) for all the powder batches and it was about 2.55-2.60 g/cm3. The location- dependent PBD values were quite interesting indicating powder segregation during spread.

The location- dependent PBD values were quite interesting indicating powder segregation during spread. It is usually theorized that finer particles typically deposit near to the feedstock (i.e., front) whereas the coarser ones are pushed toward away from the feed bin (i.e., back) [5].In addition, to understand how the powder packing state and flowability can contribute to the defect content, all XCT coupons were characterized (see Fig. 5). As shown in Fig. 5, when the virgin powder was used for fabrication, the relative densities for coupons placed near the spatter- rich regions (i.e., S and SW) were at the lowest compared with other locations. In the case of virgin powder, the maximum number of defects was detected in the SW location. When the powder was reused 7 times, the relative densities across the build plate were more consistent with less variation. Overall, it seems that the number of defects can increase by going toward the gas flow outlet.

Overall, it seems that the number of defects can increase by going toward the gas flow outlet. The maximum defect size for each specific location is also illustrated in Fig. 5. It needs to be specified that the maximum defect size was evaluated based on the assumption that all defects are spherical. As a result, the equivalent sphere diameter was used to compare the defect sizes.1099(a)(b)Fig. 5 Relative density of the XCT coupons in different locations for (a) virgin-print set and (b) 7x reused-print sets. The maximum defect size in each location is labeled on each box in microns.The maximum defect size in the case of 7x-reused powder was 55 µm. This can show that the maximum defect sizes were still comparable between the first and last fabrication sets. Interestingly, a significantly large number of defects were noted in the SW locations when the virgin powder was used (almost more than 1200 defects per cm3).

Interestingly, a significantly large number of defects were noted in the SW locations when the virgin powder was used (almost more than 1200 defects per cm3). Observing the maximum number of defects in this region can be attributed to some spatters landing in this zone as reported by Ref [17]. The high amount of defects in this area may verify that spatters can deteriorate the fusion and bonding between the subsequent layers as they typically have less absorptivity because of the oxide films around them.To provide an initial understanding of whether the variation in powder characteristics resulted from the powder reuse practice as well as the build location can affect the mechanical properties, the Rockwell C hardness (HRC) value was evaluated for each condition. As seen in Fig. 6, although negligible, it seems that the HRC is increasing with continuously reusing the powder.

As seen in Fig. 6, although negligible, it seems that the HRC is increasing with continuously reusing the powder. The slightly higher HRC values of the Print 8 specimens can be possibly due to the increase in oxygen concentration with powder reuse practice which is common in Ti64 [13]. The observation on hardness was consistent with the results reported in Ref. [29] where a higher hardness was reported for the L-PBF specimens fabricated with 12-times reused powder compared with the virgin one. This increase in hardness was correlated with the oxidation of some Ti64 particles during the sieving process or removal of the parts [29]. However, thorough analyses on chemical compositions are needed to better justify this observation. In terms of location, no specific trend was noticed. Observing no correlation between the build location and HRC can be ascribed to less sensitivity of hardness tests to defect content, which is mainly affected by the powder spreadability.1100Fig.

Observing no correlation between the build location and HRC can be ascribed to less sensitivity of hardness tests to defect content, which is mainly affected by the powder spreadability.1100Fig. 6 The Rockwell C hardness (HRC) values for different powder reuse iterations in different locations.ConclusionsThis study investigated the effects of powder reuse as well as the build location on the powder characteristics, PBDs, defect content, and HRC values. The variations in results were explained by the changes in powder characteristics resulted from the powder reuse practice and location.

The variations in results were explained by the changes in powder characteristics resulted from the powder reuse practice and location. The following conclusions can be drawn:Continuously reusing the powder resulted in a better packing state (~ 30%), flowability (~ 70%), and increased mechanical interlocking (~ 12%).Part density slightly increased with reusing the powder which was explained by the improved packing behavior of reused powder.When the virgin powder was used for fabrication, a large number of defects (more than 1200 1/cm3) were noted in the SW location which was close to the inert gas outlet. This observation was ascribed to the spattering during fabrication.HRC values showed a slightly increasing trend with powder reuse (~ 2%) which was attributed to the possible oxygen pickup during fabrication. However, no trend was uncovered with the location of the parts.

fects of Powder Variation on the Microstructure and Tensile Strength of Ti6Al4V Parts Fabricated by Selective Laser MeltingHengfeng Gu*, Haijun Gong*, J. J. S. Dilip*, Deepankar Pal*, Adam Hicks†, Heather Doak†, Brent Stucker*Department of Industrial Engineering, J. B. Speed School of Engineering, University of Louisville, Louisville, KY 40292 †Mound Laser & Photonics Center, Dayton, OHAbstractMetallic powders are used as raw materials in the Selective Laser Melting (SLM) process. These metal powders are typically available from more than one powder vendor. Even when powders have the same nominal chemical compositions, powders produced by different companies typically result in different powder particle size distributions and morphologies. These powder differences result in different powder bed thermophysical properties, which affect how the powder melts and solidifies.

These powder differences result in different powder bed thermophysical properties, which affect how the powder melts and solidifies. This paper studies the effect of powder variation on the microstructure and tensile strength of as-built SLM Ti6Al4V parts. Ti6Al4V powders from different vendors were used to fabricate parts via SLM. Powder characteristics, such as particle size distribution, morphology, and flowability, were obtained. Powder bed densities and thermal conductivities were measured and compared. The microstructures and tensile strengths were investigated by standard metallographic and tensile testing methods.

Powder bed densities and thermal conductivities were measured and compared. The microstructures and tensile strengths were investigated by standard metallographic and tensile testing methods. Based on the experimental results, a correlation between the powder characteristics and part properties are discussed.IntroductionSelective laser melting (SLM) is a powder bed fusion based additive manufacturing process, which utilizes a laser beam as the energy source to scan and melt selected areas of metallic powders in order to fabricate a 3D part in a layer-by-layer fashion. The Ti-6Al-4V alloy is widely used in SLM for aerospace, automotive and biomedical applications due to its unique characteristics such as high strength to weight ratio, excellent corrosion resistance and wear performance [1,2,3].Metal powders used in SLM are commercially available from different vendors.

Due to different techniques and specifications used in powder manufacturing, different powder lots can have different properties even if they have the same chemical compositions. Such differences in powder properties may affect the powder bed thermophysical properties, which may later affect the as- built SLM part properties.Liu et al investigated 316L stainless steel powders with 2 types of particle size distributions and the properties of as-built part using SLM, of which the results indicate that powders with different particle size distributions behave differently and thus cause a difference in an as-built part’s quality [4]. Spierings et al found that a difference in the particle size distributions of 316L stainless steel powders from 2 different vendors affect as-built part density, surface quality and mechanicalproperties [5].

Few studies have been done to investigate the relationship between the properties of powders, the powder bed and as-built parts in a comprehensive way.This paper investigates the effect of Ti-6Al-4V powder variation on the powder bed thermophysical properties and the microstructure and tensile strength of as-built SLM parts. A model-guided processing parameter set is also applied in the SLM fabrication process to achieve nearly identical as-built part properties when Ti-6Al-4V powder properties vary.Materials and experimental proceduresMaterial selection and characterizationTitanium pre-alloyed powders acquired from 3 different suppliers (EOS GmbH, LPW Inc. and Raymor Industries Inc.) were used in this study. The three types of powders with identical nominal chemical composition of 6wt% aluminum and 4wt% vanadium were manufactured using gas atomization. These 3 powder types were characterized according to their particle size distribution, morphology and flowability.

These 3 powder types were characterized according to their particle size distribution, morphology and flowability. Particle size distribution was captured using a particle analyzer (Microtrac S3000). The morphology of each type of powder was observed under a scanning electron microscope (FEI Nova Nano SEM). Angle of repose (AOR) was chosen as an indicator of powder flowability, which was obtained using a 3 inch diameter sectioned PVC pipe with a clear end film. The largest angles of the powder before it slipped during gentle counter-clockwise manual rotation of the pipe are shown in figure 1.

The largest angles of the powder before it slipped during gentle counter-clockwise manual rotation of the pipe are shown in figure 1. Three replicates were done for each type of powder.Counter- clockwise RotationPowder Slip PlaneAORInitial Horizontal PlaneFigure 1 Angle of Repose (AOR) experimentation for EOS Ti6Al4V powderPowder bed characterizationPowder bed density and thermal conductivity are two factors which influence the Selective Laser Melting process since they may dramatically influence powder solidification and consolidation, thus impacting the quality of an as-built part. [4,6] Two simple preliminary experiments were designed and implemented in order to measure the values of these two factors for each type of Ti6Al4V powder bed.A specially designed laser-milled stainless steel disk of 60mm diameter was used to measure the powder bed density of each type of Ti6Al4V powder.

The disk was put on the build plate of the Direct Metal Laser Sintering Process Development Cell (DMLS PDC) at Mound Laser & Photonic Center, Inc (MLPC) and three layers of 0.03mm thick powders were spread over it for each measurement, which created a total 0.09mm height. The volume of the powder therefore could be determined. The disk was then removed from the platter and weighed both with and without powders, of which the difference is the mass of 3 layers of powders. The powder bed density was calculated using equation (1), where 𝜌𝑃𝑜𝑤𝑑𝑒𝑟−𝑏𝑒𝑑 is the powder bed density, 𝑚𝑃𝑜𝑤𝑑𝑒𝑟−𝑏𝑒𝑑 is the mass of powders on the disk, and 𝑉𝑃𝑜𝑤𝑑𝑒𝑟−𝑏𝑒𝑑 is the volume of powders.

The powder bed density was calculated using equation (1), where 𝜌𝑃𝑜𝑤𝑑𝑒𝑟−𝑏𝑒𝑑 is the powder bed density, 𝑚𝑃𝑜𝑤𝑑𝑒𝑟−𝑏𝑒𝑑 is the mass of powders on the disk, and 𝑉𝑃𝑜𝑤𝑑𝑒𝑟−𝑏𝑒𝑑 is the volume of powders. 3 replicates were made for each powder type.𝜌𝑃𝑜𝑤𝑑𝑒𝑟−𝑏𝑒𝑑 =𝑚𝑃𝑜𝑤𝑑𝑒𝑟−𝑏𝑒𝑑 𝑉𝑃𝑜𝑤𝑑𝑒𝑟−𝑏𝑒𝑑In order to measure the powder bed thermal conductivity, a series of steady state thermal profiles were generated in the MLPC DMLS PDC using a range of fiber laser power intensities and a 125- 250 µm diameter laser beam so that the powder bed was gradually heated. The area covered by the illuminated laser is in the form of a circle with diameters of 1.25, 1.75 and 2.54 mm, where the hatch spacing between consecutive laser scans in the layer has been maintained at 100 µm with a laser scan rate between 1000-5000mm/s to avoid powder melting. The processing parameters for powder bed thermal conductivity measurements are shown in table 1.

The processing parameters for powder bed thermal conductivity measurements are shown in table 1. The scan was continued for a minute in order for the powder bed to achieve a steady state thermal distribution. The schematic for the gradual laser heating scan pattern approximated using equation (2) is shown in figure 2, where 𝑅 is the radius of the laser illuminated circle, ℎ is the hatch spacing between consecutive laser scans and 𝑥 is the x-axis coordinate for a specified spot within the laser illuminated circle and 𝑦 is its y-axis coordinate.

A FLIR camera was mounted beside the DMLS PDC for in-situ thermal profile recording, and the obtained thermal data were later analyzed for thermal conductivity calculations using the derived equation (3), where 𝑘 is the calculated thermalconductivity value, 𝐴 is the absorptivity of the powder, 𝑃 is the laser power, (𝜌𝑝𝑜𝑤𝑑𝑒𝑟−𝑏𝑒𝑑 𝜌𝑏𝑢𝑙𝑘fraction density of the powder bed with respect to the Ti6Al4V bulk density, 𝛼 is the powder absorptivity rate, 𝑟 is the distance between a specified spot in the laser illuminated circle and(1)) is thegeometrical center of the circle, 𝑅 is the radius of the laser illuminated circle, ∆𝑇 is the temperature difference between the steady state temperature at the specified spot with a distance of 𝑟 to the2𝑅geometrical center of the laser illuminated circle and room temperature, and 𝑣𝑥 = is the time for finishing the illumination of the circle once (computed using a calculation of the, in which T𝑇approximate arc length divided by the scan speed= 4𝑛 𝑖=1∑ (√(𝑅2−(𝑖ℎ)2))𝑣where 𝑖 is number ofintercepts on the x axis, ℎ is hatch spacing and 𝑣 is scan speed).Table 1 Processing parameters of Ti6Al4V powder bed thermal conductivity measurement experimentParameters Beam diameter (μm) Scan speed (mm/s) Hatch spacing (μm) Laser power (W) Scan circle diameter (mm)Values 125, 175, 250 1000, 2000, 3000, 4000, 5000 100 45, 68, 90 1.25, 1.75, 2.54𝑦 = √(𝑅2 − 𝑥2)sin (𝜋𝑥 ℎ)(2)10.5 1.5x 10-31.50.5 x axis (in meters)y axis (in meters)11.5Figure 2 Schematic of the gradual laser heating scan pattern utilized for the measurement of Ti6Al4V powder bed thermal conductivity𝑘 = 𝐴𝑃 (𝜌𝑝𝑜𝑤𝑑𝑒𝑟−𝑏𝑒𝑑 𝜌𝑏𝑢𝑙𝑘) (1 − 𝑒−𝑣𝑥𝛼𝑟 4𝜋𝑟∆𝑇𝑅𝑣𝑥𝛼)Testing specimen fabrication processTensile testing specimens were designed according to ASTM E8/E8M-1.

They were fabricated using EOS, LPW and Raymor Ti6Al4V powders respectively using an EOS DMLS M270 machine in argon purging atmosphere at the University of Louisville. An extra 5 mm was appended on both sides of the tensile testing specimen in the grip section for further microstructural analysis. Two sets of tensile testing specimens were fabricated as follows and 3 replicates of each parameter set were made:Case A: Specimens were fabricated using identical parameters for EOS, LPW and Raymor Ti6Al4V powders respectively in order to investigate the effect of powder variation on the as-built part quality under identical fabrication conditions.

This parameter set was obtained by performing an empirically optimization of parameters for Raymor Ti6Al4V powders, as shown in table 2.Case B: EOS and LPW Ti6Al4V specimens were fabricated using model-guided processing parameter sets based on Case A by changing the hatch spacing values to keep a constant melt pool overlap for all 3 types of powders, as shown in table 3. The modified processing parameter sets were developed using novel finite element software developed at the University of Louisville [8].

The modified processing parameter sets were developed using novel finite element software developed at the University of Louisville [8]. The hatch spacing used in the machine was rounded up from the model predicted value to the next highest multiple of 10, as EOS software only accepts hatch spacing inputs as multiples of 10.Table 2 Identical processing parameters for EOS, LPW and Raymor Ti6Al4V powders (Case A)Laser Power (W) Scan Speed (mm/s) Layer Thickness (μm) Hatch Spacing (μm)186120030100(3)Table 3 Model-guided processing parameters for EOS and LPW Ti6Al4V powders based on Raymor empirically optimal parameter (Case B)Supplier Laser Power (W) Scan Speed (mm/s) Layer Thickness (μm) Hatch Spacing (μm)EOS LPW186 1861200 120030 3080 90Raymor186120030100Microstructural analysisThe extra grip section material was removed from the tensile testing specimen using a Buehler Isomet slow speed saw.

These materials were further prepared using standard metallographic specimen preparation methods. Microstructures in YZ planes of 5 specimens were observed using an optical microscope (Olympus MX51).Tensile testingThe support structure was milled off the bottom surface of each tensile testing specimen, but all other surfaces were kept in their as-built condition. Tensile tests were carried out using an INSTRON system at the University of Louisville. An extensometer was used to record strain-stress data during the testing and 3 replicates were done for each processing parameter set.Results and DiscussionsMaterial characterization(a) Particle size distributionAccording to figure 3, EOS and LPW Ti6Al4V powder particles are uni-modally distributed within the range from 20μm to 50μm, while Raymor Ti6Al4V powder particles have a bi-modal distribution with a major particle size range from 20μm to 50μm and a minor particle size range from 4μm to 10μm.

The mean volume diameter value (MV) and its corresponding standard deviation (SD) for the 3 types of powders are shown in figure 3.MV (μm)SD (μm)38.0410.18MV (μm)SD (μm)33.218.71MV (μm)SD (μm)29.949.7Figure 3 Particle size distribution graphs for EOS, LPW and Raymor Ti6Al4V powder particles(b) Particle morphologyFigure 4 shows SEM images of EOS, LPW and Raymor Ti6Al4V powders. A majority of particles in all 3 types of powders appear spherical in shape. In particular, LPW powders seem to be more spherical compared to the other two types of powders, while a noticeable amount of particles with irregularly deformed shapes were found in EOS and Raymor powders according to the SEM images.

In addition, there exists a significantly large amount of agglomerated fine particles in Raymor powders compared to EOS and LPW powders due to the bi-modal distribution in Raymor powder particles, which matches the PSA results.Figure 4 SEM images for EOS, LPW and Raymor Ti6Al4V powders(c) Powder flowabilityAngle of repose (AOR) is considered a good indicator of powder flowability; better powder flowability can be achieved at a smaller AOR value [7]. The comparison of AOR values for each type of powders is shown in figure 5, where a large difference of AOR values in 3 types of powders was found. The AOR value of Raymor powder is significantly higher than the other two types of powders, indicating that Raymor powder has the poorest flowability. This is probably due to the fine agglomerates within the powder, which increases inter-particle friction and reduces powder flowability.

This is probably due to the fine agglomerates within the powder, which increases inter-particle friction and reduces powder flowability. Meanwhile, AOR value for LPW powder is the lowest among the 3 types of powders, which may be explained by its noticeably more spherical particle shape found from the SEM images.Figure 5 Angle of repose values for Ti6Al4V powders from 3 suppliersPowder bed characterization(a) Powder bed densityFigure 6 shows the comparison of Ti6Al4V powder bed density (PBD) values from 3 suppliers. There is no significant difference in PBD values for 3 types of powders; the PBDs of all 3 powders are approximately 47% of the Ti6Al4V bulk density of 4.42g/cc.

There is no significant difference in PBD values for 3 types of powders; the PBDs of all 3 powders are approximately 47% of the Ti6Al4V bulk density of 4.42g/cc. The results indicates that powder flowability may not have a crucial impact on the powder bed density, even though it is reported to have an influence on the particle distribution in the powder bed [4].Figure 6 Ti6Al4V powder bed density from 3 suppliers(b) Powder bed thermal conductivityFigure 7 shows the steady state thermal profiles of the powder beds for 3 types of Ti6Al4V powders in the laser illuminated areas. Raymor powder achieves much lower temperatures than EOS and LPW powder beds, indicating that the thermal properties of Raymor powder in the bed may be significantly different from EOS and LPW powder beds.

One possible explanation can be that a lot of agglomerated fine particles within the Raymor powders may be easily sintered to the adjacent large powder particles at a relatively lower temperature, which may dramatically increase the overall thermal conductivity of the powder bed. The approximate percentage of the thermal conductivity for the powder bed compared to bulk Ti6Al4V properties were obtained using equation (3), and are shown in table 4. The trend of the calculated thermal conductivity values of 3 types of powders matches the above explanation.Figure 7 Steady state thermal profiles of (a) EOS, (b) LPW, and (c) Raymor Ti6Al4V powder bed using FLIR camera.

(Images are not shown at the same scale.)Table 4 Approximate percentage of thermal conductivity for the powder bed compared to bulk Ti6Al4VSupplier EOS LPW RaymorApproximate percentage of thermal conductivity of bulk Ti6Al4V (%) ~20 ~20 ~60Microstructural analysis for as-built test specimensFigure 8 shows the typical microstructures of specimens fabricated using the identical Raymor- based empirically optimal parameter set (Case A), where α’ martensite phases dominate the entire microstructure in all specimens. Due to the high cooling rate, all specimens experienced similar phase transformations during fabrication resulting in similar microstructures. α’ columnar grains with prior β grain boundaries grow in the build direction in the YZ plane. However, serious porosity occurs in specimens which were fabricated using EOS and LPW Ti6Al4V powders, while Raymor specimens achieve near full density under identical processing parameters.

However, serious porosity occurs in specimens which were fabricated using EOS and LPW Ti6Al4V powders, while Raymor specimens achieve near full density under identical processing parameters. The difference in as-built part porosity is possibly linked to the disparity of the powder bed thermal properties during the fabrication process. When the same amount of laser energy is applied, more energy will be transmitted into the surrounding Raymor powder and thus form a wider melt pool and larger overlap than those formed using EOS and LPW powders due to the higher thermal conductivity in the Raymor powder bed.

When adjacent melt pools do not overlap with each other, pores started to form in EOS and LPW specimens.Figure 8 Microstructures in YZ plane of specimens fabricated using (a) EOS, (b) LPW, and (c) Raymor Ti6Al4V powders for identical Raymor-based empirically optimal parametersThe model-guided processing parameter sets for EOS and LPW powders (Case B) were applied during fabrication by only narrowing hatching spacing in order to achieve the same melt pool overlap as Raymor’s empirical optimal parameter set, thus seeking to avoid porosity in as-built EOS and LPW Ti6Al4V specimens.

Comparison among microstructures of EOS, LPW and Raymor specimens fabricated using Case B are shown in figure 9, where all 3 specimens have reached near full density with the same phase composition as in figure 8.Figure 9 Microstructures in YZ plane of specimens fabricated using (a) EOS, (b) LPW, and (c) Raymor Ti6Al4V powders based on model-guided processing parameter setsTensile testingFigure 10 shows the UTS and yield strength values for tensile specimens fabricated using Case A processing parameters. Both UTS and yield strength values for all 3 types of specimens were around the same level, which are approximately 1200MPa and 1100MPa respectively.

Both UTS and yield strength values for all 3 types of specimens were around the same level, which are approximately 1200MPa and 1100MPa respectively. The results indicate that the powder type variation does not have a significant influence on the as-built part’s tensile strength, even though it may dramatically affect the part’s fatigue performance due to the large difference in porosity.Figure 10 UTS and yield strength for parts fabricated using EOS, LPW and Raymor Ti6Al4V powders under Raymor-based empirically optimal processing parametersFigure 11 shows the comparison of UTS and yield strength values for tensile testing specimens fabricated for Case B. Similarly to figure 10, no significant difference was found in UTS and yield strength values for parts manufactured using powders from 3 different suppliers.

Similarly to figure 10, no significant difference was found in UTS and yield strength values for parts manufactured using powders from 3 different suppliers. In addition, both UTS and yield strength values are at the same level as those in figure 10.Figure 11 UTS and yield strength for parts fabricated using EOS, LPW and Raymor Ti6Al4V powders using model-guided processing parameter setsSummaryTi-6Al-4V pre-alloyed powders obtained from 3 different suppliers were investigated in this study. Powder properties including particle size distribution, morphology, and flowability were characterized using various methodologies. All 3 types of powders appear spherical in shape,however, a significantly large amount of fine particles were found in Raymor powders compared to the other two types of powders due to its bi-modal distribution.

The flowability of Raymor powder is also much poorer than the other two powders caused by inter-particle friction introduced by these fine particles.Powder bed densities of 3 types of Ti-6Al-4V powders were determined and no significant differences were found.

However, the calculated Raymor powder bed thermal conductivity value was significantly higher compared to the values of the other two powder beds based on the powder bed thermal profiles captured using the thermal camera, which indicates that the thermal properties of Raymor powder may be completely different from the ones of the other two powders.Tensile and metallurgical specimens were fabricated successfully under two sets of processing parameters; empirically optimal Raymor processing parameters for all 3 types of powders and model-guided processing parameters for each type of powder.Serious porosity were found in EOS and LPW specimens using empirically optimal Raymor processing parameters, while the Raymor specimen was near fully dense. Specimens fabricated using all 3 types of powders under model-guided processing parameter sets achieved near full density.

Specimens fabricated using all 3 types of powders under model-guided processing parameter sets achieved near full density. No significant difference in tensile strength were found in specimens fabricated using all 3 types of powders under all processing parameter sets.Microstructures in all specimens appear very similar, where α’ columnar grains with prior β grain boundaries grow in the build direction in the YZ plane.Future workThe accuracy of the powder bed thermal conductivity method used in this study has not been determined. As it appears that powder bed thermal conductivity may be a significant factor for process parameter optimization, a well-established experimental for these types of experiments should be developed and its accuracy determined.

Additionally, since there is no conclusive explanation for Raymor’s significantly higher thermal conductivity value compared to the other two types of powders, further studies are needed to investigate the mechanism that leads to this large disparity.No significant difference was found in tensile strength values between specimens with and without porosity, which was an unexpected outcome. More replicates are needed to make sure the validity of the effect of porosity on the as-built part tensile properties are valid.AcknowledgementThe authors thank the staff of the Rapid Prototyping Center at the University of Louisville and Mound Laser & Photonics Center, Inc for their help. We gratefully acknowledge the support of the Air Force Research Laboratory through contract FA8650-13-M-5060.References[1] Antonysamy, A. A. (2012).

E EFFECTS OF PRE-ALLOYED STEELS POWDER COMPOSITIONS ON BUILD QUALITY IN DIRECT METAL LASER RE-MELTING.SP Akhtar, CS Wright and M Youseffi Engineering Materials Research Group, University of Bradford, Bradford, UKAbstractThis paper describes recent studies investigating the relationships between alloy composition and processability of a range of pre-alloyed stainless and tool steel powders melted using CO2 and Nd:YAG lasers. Factors considered in this paper that influence processability are particle size and shape, particle distribution, fluidity, wetting, and scan conditions. However, thermodynamic modelling, validated by experiment, suggests that alloy composition is a significant factor affecting processability. A short solidification range corresponds to a significant increase in build quality.IntroductionSelective laser sintering (SLS) is one of the more important solid freeformmanufacturing processes developed in the last 10-15 years [1].

One of the main applications of SLS is for the manufacture of prototype or low volume production tooling using steel powders. Commercial systems are available which generate parts from either polymer coated steel powders [DTM Rapid Tool 2.0TM] or by processing a mixture of metal powders that contains a low melting point component [EOS Direct Tool TM]. The need to post process, i.e. infiltrate with bronze, or the reliance on a low melting point liquid, leads to components with poor mechanical and tribolological properties. This has led to a number of studies aimed at investigating Direct Metal Laser Re-melting (DMLR) pre-alloyed metal powders with the aim of eliminating the above disadvantages. These studies generally made use of commercially available powder stainless steel [2, 3], and tool steels [4 - 6]. Stainless Steel powders have been DMLR processed to give multi-layer samples with densities of +99% [3] compared with only 60% for HSS [6].

Stainless Steel powders have been DMLR processed to give multi-layer samples with densities of +99% [3] compared with only 60% for HSS [6]. This paper presents some results of ongoing investigations into the melting / densification behaviour of pre-alloyed metal powders during DMLR which has the ultimate aim of identifying the factors which control the ability to fabricate fully dense, multi- layer components.Experimental ProcedureFive types of powder were used in this investigation. Annealed water atomised M2(WA-M2), gas atomised M2 (GA-M2), gas atomised H13 (GA-H13) with varying contents of carbon and water atomised P58 (WA-P58), an experimental high speed steel (Fe-14Mo-4Cr- 14C) . The water atomised M2 was sourced from Powdrex, UK and the gas atomised powders were obtained from Osprey Metals Ltd, UK. Water atomised powders were sieved in compliance with ASTM standard E11 to give the following size fractions; >75-150µm, >38- 75µm and <38µm for water atomised powders.

Water atomised powders were sieved in compliance with ASTM standard E11 to give the following size fractions; >75-150µm, >38- 75µm and <38µm for water atomised powders. Identical size fractions were obtained for gas atomised powders from the suppliers.Direct metal laser re-melting studies were performed on specially constructed machines at the Universities of Leeds and Liverpool. The Leeds machine includes a 250W continuous wave (CW) CO2 laser. Initial studies were conducted at a beam size of 1.1mm diameter. This was subsequently reduced to 0.55mm diameter. Sintering was performed in an141argon shroud. Details of this machine are given in Reference [2]. The Liverpool system, described by Morgan, et. al. [3], consisted of a Rofin Sinar 90W, flash lamp pumped Q- Switched Nd : YAG laser. An analogue galvanometer-scanning head attains beam position over an area of 80 x 80 mm2. Pulse repletion rates were in the range 0 – 60kHz and minimum beam diameter is 80µm.

An analogue galvanometer-scanning head attains beam position over an area of 80 x 80 mm2. Pulse repletion rates were in the range 0 – 60kHz and minimum beam diameter is 80µm. The shroud gas was nitrogen.The two systems employed different approaches to placing layers of powder. In theLeeds machine, the initial powder layer comprised loose powder to a depth of 5 mm contained in a mild steel tray. This layer was levelled with a blade to ensure a flat powder surface. Subsequent layers were deposited using a hopper system. The layer depth was 0.4 mm. In the Liverpool machine mild steel sheet was used as a substrate. In this system, the powder delivery system was optimised for an even 100µm coating of powder to be layered for every build layer [3].

In this system, the powder delivery system was optimised for an even 100µm coating of powder to be layered for every build layer [3]. Whilst all size fractions studied could be processed on the Leeds machine, the Liverpool machine was restricted to processing just the <38µm fraction.After sintering samples were prepared for metallographic examination by mountingand grinding on SiC paper to 1200 grit, then polished on cloths impregnated with 6 and 1µm diamond. Polishing was completed using 0.05µm γ-alumina. Samples were studied in the etched and un-etched conditions. Samples were etched in 5 or 10% Nital. Samples were examined using both optical and scanning electron microscopy techniques.Isopleths were calculated for the M2, H13,316L, P58 systems using ThermoCalcTM[www.thermocalc.se] software and the SGTE SSOL database in order to understand the correlation between alloy composition, solidification ranges and processability.Results and Discussions.Powders.

Typical SEM micrographs for the different powders are given in Figures 1 and 2.Fig1 . M2 water atomised HSS powder. As received <150µm. SEM X150, 20KVFig 2. P58 gas atomised HSS powder, As received. SEM X1000, 20KVThe water-atomised powders were composed of highly irregular particles and, as such, are typical of this method of powder production. As expected the gas-atomised powders were composed of spherical particles. However, whilst the <38µm GA-P58 size fraction consisted of spherical particles with smooth surfaces the particle surfaces of the coarser fractions were much rougher due to the presence of many smaller satellites formed as a result of collisions during atomisation, Figure 2. The effects of powder size and morphology on DMLR are discussed later.

The effects of powder size and morphology on DMLR are discussed later. The various size fractions used, along with Apparent Density and flowability (Hall Flowmeter) data are given in Tables 1 and 2.142Table 1 Powder properties for WA-M2 HSS and WA-P58 powder.Powder Fraction (µm)150 +75/-150 +38/-75 -38WA-M2 HSS powder Flow rate (s/50g) 52.1 55.4 59.8 No flowWA-M2 HSS powder Apparent Density (g/cm3) 2.27 2.14 2.18 2.60WA-P58 Apparent Density (g/cm3) 2.70 2.20 2.52 2.87WA-P58 Hall Flow rate (s/ 50g)No flow 51.3 No flow No flowTable 2 Powder properties for GA-M2 HSS and GA-P58 powderPowder Fraction (µm) Apparent Density (g/cm3)GA-P58 (As-Received –150) GA-M2 -75/+150 GA-M2 -385.003.974.31Hall Flow rate (s/ 50g) 18.223.7No flowAt the outset of this work two methods were envisaged for the manufacture of highdensity, multi-layer components as discussed in previous work [7].

The preliminary strategy was to partially melt successive powder layers together followed by a post processing step. The second method was to completely re-melt the powder layers. In order to identify appropriate processing conditions for the two processing strategies, the heating and melting behaviour of each batch of powder was systematically studied using the CO2 laser SLS machine at Leeds. Single tracks were produced for a wide range of Laser power and scanning conditions.Process Maps. Single track process maps were produced using the CW CO2 DMLR machine for M2, H13 and P58. The process maps detail the heating and melting behaviour of the metal powders as scan speed and laser power change. These highlight regions where the melt pool remained continuous, i.e. did not fragment or ball. A fragmented melt pool has been shown to reduce the surface quality and density of layers [8]. Typical Process Maps for M2 and H13 powders are given Figure 3.

did not fragment or ball. A fragmented melt pool has been shown to reduce the surface quality and density of layers [8]. Typical Process Maps for M2 and H13 powders are given Figure 3 . The maps show several qualitative regions: continuous flattened tracks, continuous rounded tracks, melted and balled tracks, partially melted and unmelted regions.143(a)(b)Figure 3. CW CO2 laser process maps for (a) gas-atomised M2 <150/+75µmfraction 1.1mm spot size (b) gas-atomised H13 <150/+75 µm fraction, 0.6mm spot sizeA feature of these process maps for single track scanning of M2 and H13 powders wasthat the size of the regimes and the positions of the boundaries defining each regime were comparable, irrespective of alloy composition, or whether or not the powder was gas or water atomised, or processing with spot sizes of 1.1 or 0.6mm diameter, or even size fraction used, e.g. Figures 3(a) and (b).

Figures 3(a) and (b). Moreover, the maps were comparable to ones reported elsewhere for 314HC stainless steel [8] and Fe-14Mo-4Cr-1.4C [9] experimental HSS alloys.In order to fabricate multi-layer blocks (12 x 12 x 12 mm3), experiments were performed on the CO2 laser SLS machine using conditions within both the partial melting and continuous melting regimes. However, the sub 38µm water and gas atomised size fractions were difficult to spread due to their poor flowability, layers tended to form furrows. Work with these powders was, therefore, discontinued. With the coarser fractions, irrespective of powder type or particle sizes, processing within the partial melting regime was unsuccessful because insufficient melting took place to bond the next layer on to the previous layer, making it impossible to build intact blocks [10]. Thus it was not possible to investigate fabrication of multi-layer components for M2 via the partial melting-post processing route.

Thus it was not possible to investigate fabrication of multi-layer components for M2 via the partial melting-post processing route. Gas atomised M2 multi-layer blocks fabricated from 25 individual powder layers were produced using conditions within the continuous melting regime. The resultant blocks were very porous and had very rough surfaces. Typical micrographs consisted of M2C and M6C eutectics dispersed in a martensitic matrix and a definite interfacial feature was also present between bonded layers comprising eutectic carbide structures [7].On processing the +75/-150µm H13 size fraction in the CW CO2 system it was noted that for scan conditions in the low melt aspect ratio regime, Figure 3 (b), produced tracks that144were 30-40% flatter than for M2 powders. So from a build point of view H13 had better processability in the CW CO2 laser system but again multi-layer blocks produced from H13 powders were very porous and fragile, and only a maximum density of 70% was obtained.

As with M2, processing H13 powders in the partial melting regime was unsuccessful since it was not possible to bond the melted layer to the underlying layer.Much more promising results were obtained with the Nd : YAG DMLR machine. Thissystem employed much higher scan speeds and a layer thickness of 0.1mm. In the Nd : YAG system no systematic process maps were collated. Conditions were selected based on processing parameters used to produce high density multi-layer blocks from 316L powders [3]. Water atomised M2 powders were processed at scan speeds of 10 –500mms-1, laser current 16-20A, Spot size = 0.1 mm, Beam overlap of 25-80%, pulse repetition frequency of 0kHz, layer thickness 100µm. Both single and 3 layer samples were produced but the scanned areas contained large amounts of un-melted powder. Removal of this by ultrasonic cleaning revealed that the beds were very porous. The porosity levels were unaffected by scanning conditions.

Removal of this by ultrasonic cleaning revealed that the beds were very porous. The porosity levels were unaffected by scanning conditions. This high level of porosity was attributed to the low apparent density of the water atomised powder (Table 1) arising from its highly irregular shape.Further experiments with WA and GA M2 powders were carried out at layer thicknesses of 100 and 50µm, pulse repetition frequencies of 40 and 50kHz, scan speed = 500 mms-1, current = 20A, spot size = 100µm, beam overlap = 50%. Blocks 5 x 15 x 15mm3 were produced. Operating in pulse repetition mode resulted in an increase in density compared to continuous mode. However resultant densities, which were the highest attained in this study on laser scanning of M2 powders, were only 5.5 gcm-3, i.e. ~67% relative, Figure (a).

However resultant densities, which were the highest attained in this study on laser scanning of M2 powders, were only 5.5 gcm-3, i.e. ~67% relative, Figure (a). Despite the higher apparent density of the <38µm gas atomised powder, Table 2, use of this powder had no beneficial effects on the density of laser scanned blocks, densities were similar to those obtained for <38µm water atomised M2 powder. So it would appear that in pulsed mode final density is independent of initial powder bed density. Also for the range of conditions studied it would appear that density is unaffected by pulse frequency.(a)(b)Figure 4. Micrographs of Nd : YAG laser scanned multi-layer (a) GA-M2, 500mms-1scan speed, 50% beam overlap, current 20A, spot size 0.1mm, pulse rate 40kHz. (100µm layer thickness). (b) GA-H13 <22µm.

Micrographs of Nd : YAG laser scanned multi-layer (a) GA-M2, 500mms-1scan speed, 50% beam overlap, current 20A, spot size 0.1mm, pulse rate 40kHz. (100µm layer thickness). (b) GA-H13 <22µm. Scan speed = 500mms-1, beam overlap = 50%, laser current = 20A, spot = 0.1mm, pulse rate 30kHz.145Our work indicates that since the maximum density for M2 was only 60% relative, and 67% relative using CW CO2 and Nd : YAG systems, respectively this is not a suitable alloy for processing by DMLR. Moreover the microstructures produced contained detrimental microstructural features such as carbide clusters and thermal cracking [7]. Consequently, even if M2 powders could be processed to ~100% density by DMLR, such regions would also be potential points of weakness. Carbide clusters are preferred sites for crack nucleation and propagation at low applied stresses under static and dynamic loading [11].

Carbide clusters are preferred sites for crack nucleation and propagation at low applied stresses under static and dynamic loading [11]. Given that solidification behaviour of individual HSS compositions are all similar, it is very likely that similar features would occur with other types of HSS processed by DMLR, suggesting that HSSs are unsuitable for processing by this fabrication route.However, processing GA-H13 at conditions of pulse repetition frequency of 30kHz gave well-bonded, multi-layer blocks with a density of 85-90% Fig 4(b). Processing under these conditions takes place within the melting with balling region of Figure 3 (b), giving a sample structure somewhat different from that achieved in the CW CO2 laser. The microstructure comprised martensitic cellular grains. Grain boundaries were delineated by continuous carbide network and the matrix was martensitic. The microstructure of Nd : YAG processed material was finer than the CW CO2 processed H13.

Grain boundaries were delineated by continuous carbide network and the matrix was martensitic. The microstructure of Nd : YAG processed material was finer than the CW CO2 processed H13. The finer microstructure is indicative of a faster cooling rate . However, thermal cracking was also present in these samples [14].In the Nd:YAG system further experiments were carried out, in which additionalcarbon was added to H13 powders [12]. It was thought that a reduction in the melting point would be considered advantageous when the material was laser sintered and hence aid the ease of processability. However it was found that as carbon is added to H13, the increased carbon content lead to severe balling of the molten metal and caused a lack of wetting with the resulting material became increasingly porous. Thus, it appears that the higher carbon levels are detrimental to the DMLR processing of tool steels, as carbon appears to reduce the flow of the melt as it wets the underlying material.

Thus, it appears that the higher carbon levels are detrimental to the DMLR processing of tool steels, as carbon appears to reduce the flow of the melt as it wets the underlying material. Higher carbon contents also increased the surface roughness [12].Given that the powder properties (flowability, AD and morphology) of gas atomisedM2 and H13 used in these studies were comparable, this would suggest that other factors have a very important role in determining the densification behaviour during the laser scanning of powders. In previous studies factors considered important on processing powder via DLMR include particle size, shape, distribution, melt viscosity, wetability, surface tension, oxide stability [3] and powder oxygen content [4,5]. For example, from the results of single scan studies Nui and Chang [4, 5] considered the high oxygen content of water atomised M2 powder compared to gas atomised M2 to be detrimental to their processability.

The results presented in the current paper suggest that alloy composition is also an important factor in determining whether high density can be achieved. This is particular so with the Nd:YAG system where it has been possible to process similar sized stainless steel powders, <22µm, to 100% density [3].In conventional supersolidus liquid phase sintering (SLPS) of pre-alloyed powders a critical factor for determining processability is the melting-freezing range since a wide range correlates to a wide process window 16]. Wright et. al. [17] have used German’s model of supersolidus liquid phase sintering [16 ], in conjunction with computer-aided alloy, to design tool steel compositions with enhanced sinterability [17]. The P58 composition was one such composition [17].146In the case of DMLR, processing generally involves solidification of a fully moltenbead rather than partially melted powders as in case of conventional SLPS.

An important factor in conventional casting is melt fluidity [18]. Melt fluidity determines the ease at which a molten alloy will fill a complex cavity. High melt fluidity is an essential requirement for a casting alloy since this facilitates the production of complex castings. High fluidities are associated with eutectic alloy compositions whereas long freezing range alloys exhibit low fluidities [18]. In the case of DMLR processing, alloy fluidity is likely to have an influence on the way in which the melt bead interacts with surrounding powder and the morphology of the resultant track as the bead solidifies.The results from thermodynamic modelling are shown in Figure 5 which presentsisopleths for M2, H13, P58 and 316L. 316L Stainless steel was modelled as a quinary Fe-Cr- Ni-2.5Mo-2Mn system since the very low carbon content of this alloy (0.03%) would not have a great effect on the solidification range.

316L Stainless steel was modelled as a quinary Fe-Cr- Ni-2.5Mo-2Mn system since the very low carbon content of this alloy (0.03%) would not have a great effect on the solidification range. Mo and Mn were included so that their ferrite and austenite stabilising effects are considered. Isopleths were calculated for Fe-17Cr-Ni- 2.5Mo-2Mn and Fe-Cr-12Ni-2.5Mo-2Mn. These are given in Figures 5(a) and (b), respectively. At the 316L composition (Fe-17Cr-12Ni-2.5Mo-2Mn) the solidus-liquidus interval is narrow, ~5oC.

Isopleth through the Fe-Cr-Ni-2Mn-2.5Mo system at 17Cr. 316LFig Fig 5b Isopleth through the Fe-Cr-Ni-2Mn-2.5Mo system at 12Ni. 316LFigure 5c Isopleth for the Fe-C-5Cr-1.5Mo-1V-1Si system. (H13)Figure 5d Isopleth for the Fe-6W-5Mo-4Cr-2V-C system after suspending the WC phase.(M2)Figure 5e. Isopleth for Fe – 14Mo – 4Cr – C calculated. P58Figure 5. Calculated Isopleths for M2, P58, 316L and H13.148For processing under comparable scan conditions, alloys with short melting / freezingranges, e.g. 316L where TLiquidus – TSolidus = ~5oC, densities up to 99% can be obtained [3] whilst long melting / freezing range alloys such as M2 (TL – TS = ~200oC) could only be processed to densities of only ~65-70% [7]. For H13 TL – TS = ~110oC and maximum density achieved is ~90%. This correlates with the degradation in processability noted for H13 powder with increasing carbon content [12].

For H13 TL – TS = ~110oC and maximum density achieved is ~90%. This correlates with the degradation in processability noted for H13 powder with increasing carbon content [12]. Thus, there would appear to be a direct correlation between alloy solidification range and powder processability. Increasing solidification range corresponds to a significant decrease in ease of processing. Hence this would imply that one of the important factors in determining a good build is melt fluidity with high fluidity being a desirable property. Verification of this would be achieved by processing pre-alloyed powders prepared from short freezing range alloys, i.e. eutectic or near eutectic compositions.Conclusions Prealloyed H13 die steel powders are more amenable to processing via DLMR than M2 high speed steel powdersThere is a correlation between processibility (resultant density and surface finish) and freezing range for high alloy steel powders processed by DMLR.

A short freezing correlates to good processibility.Acknowledgements The research reported in this paper is funded by the UK Engineering and Physical Sciences Research Council under Grant Number GR/R32222. This paper makes use of data provided by our partners C.Hauser and T.H.C. Childs (University of Leeds) and J. Xie, P. Fox (University of Liverpool). One of us (SP Akhtar) acknowledges the receipt of a Royal Academy of Engineering travel grant.References [1]KW Dalgarno & CS Wright; “Approaches to Processing Metal and Ceramics through the Laser Scanning of Powder Beds”, Society of Manufacturing Engineers, Technical Paper PE03-143, 2003. C Hauser, THC Childs & KW Dalgarno, “Selective Laser Sintering of Stainless Steel 314HC Processed Using Room Temperature Powder Beds”, proc. Solid Freeform Fabrication Symposium Austin, Texas August 9 – 11, 1999, pp273-280.

Priddy2, Linkan Bian3,41William B. Burnsed, Jr. Department of Mechanical, Aerospace, and Biomedical Engineering, University of South Alabama, Mobile, AL 36688, USA2Department of Mechanical Engineering, Mississippi State University, Mississippi, 39762, USA3Center for Advanced Vehicular Systems (CAVS), Starkville, Mississippi, MS, 39759, USA4Department of Industrial and Systems Engineering, Mississippi State University, Mississippi, 39762, USACorresponding Author, Phone: (251) 460-7706, e-mail: ayadollahi@southalabama.eduAbstractDespite appropriate planning, various incidents can stop the additive manufacturing (AM) process of metals and cause build interruption, such as power outage, lack of powder feedstock, and/or shielding gas to mention a few. Due to the layer-by-layer nature of fabrication, an interruption to the AM process can be resumed from the location where the stoppage occurred.

Due to the layer-by-layer nature of fabrication, an interruption to the AM process can be resumed from the location where the stoppage occurred. However, build interruption may adversely affect the structural integrity of the fabricated parts, by causing localized failure near the interruption location. This study aims to investigate the influence of build interruption during the laser powder bed fusion (LPBF) process on the microstructural and mechanical properties of Ti-6Al-4V and Al-Si-10Mg specimens. For the Ti-6Al-4V specimens, results indicate that tensile failures near the interruption location are most likely to happen for non-heat-treated specimens in the as-built surface condition. Whereas for the Al-Si- 10Mg specimens, the failure location is more influenced by the prolonged stoppage and air exposure.Keywords: Build interruption; Additive manufacturing (AM); Failure location; Al-Si-10Mg; Ti- 6Al-4V5921.

IntroductionIn recent years, additive manufacturing (AM) has opened new avenues for fabricating metallic parts with complex geometries [1]. Laser powder bed fusion (LPBF) is one of the most broadly applied techniques in metal AM that receives more and more attention in the industry and academia [2]. Due to the relatively low build rate of the LPBF process, fabrication of large components will be time-consuming, thus any interruption to the process during the build could result in loss of time and money. Despite appropriate planning and procedures, various incidents may stop the LPBF process, such as lack of powder quantities, power outage, faulty build data, or various probable incidents. Due to the layer-by-layer nature of the fabrication, an interruption to the LPBF process can be resumed from the location where the stoppage occurred.

Due to the layer-by-layer nature of the fabrication, an interruption to the LPBF process can be resumed from the location where the stoppage occurred. Although a continuation of the process at the stoppage location will be economic and time-saving, build interruption may adversely affect the structural integrity of the fabricated parts by causing localized failure near the interruption plane. This is due to the fact that any stop/start during the fabrication affects the thermal history, leading to changes in the microstructures and residual stresses, and consequently, the mechanical properties of the part.Despite significant research efforts focused on the effect of process parameters and build conditions [3–8] on the microstructural features and mechanical properties of LPBF parts, to date there have only been a few studies characterizing LPBF process/build interruption effects on the properties of fabricated parts [9–15]. Terrazas-Najara et al.

Terrazas-Najara et al. studied the effect of process interruption on the microstructure and tensile properties of LPBF Inconel 625, Inconel 718, and Al-Si-10Mg alloys. They applied two different protocols for a time span of one hour by keeping the coupons either in an air exposure condition or in a condition where the chamber remained closed. Their results showed that neither microstructures nor tensile properties appeared to be sensitive to process interruption [9]. Binder et al. studied the effect of a 40-minute process interruption along with air exposure on LPBF Al-Si-10Mg. They showed that almost all the interrupted samples, fabricated in a vertical direction, were broken at the interruption plane under monotonic tensile testing. They have also reported that the process interruption, accompanied by remelting, affected the tensile properties of the samples [10]. Stoll et al.

They have also reported that the process interruption, accompanied by remelting, affected the tensile properties of the samples [10]. Stoll et al. showed that a 60-minute process interruption for 316L stainless steel (SS) samples caused a reduction in tensile strength. In their work, the chamber was opened, and the powder bed was allowed to cool down to room temperature. They also reported that the location of the tensile failure is not necessarily depended on the process interruption [12], opposite to what has been reported for the LPBF Al-Si-10Mg [10,11]. Hammond et al. reported that the tensile strength and microhardness values of interrupted Al-Si-10Mg samples were 10% lower compared to the samples fabricated without interruption. In their study, the majority of the interrupted samples failed near the interruption plane under the monotonic tensile testing [11]. Richter et al.

In their study, the majority of the interrupted samples failed near the interruption plane under the monotonic tensile testing [11]. Richter et al. studied the microstructures, tensile, and fatigue properties of LPBF AlSi12 influenced by the process interruption. Their microstructural observations revealed a distinct distortion in grain morphology at the plane of interruption; however, tensile strength of interrupted and non-interrupted samples was almost similar, and the location of failure in tensile and fatigue tests did not coincide with the interruption plane [13].The literature on build interruption during the LPBF process contains studies where the findings of one study either contradict or cannot be directly compared with those of another, mainly due to the differences in the material system and the conditions of interruption (i.e., duration of the593interruption, air exposure, etc.).

Thus, more methodical research is needed to understand the effect of build interruption on the microstructural properties and mechanical performance of different alloys fabricated via LPBF process. The duration for which the LPBF machine is shut-down due to an interruption is a function of the type of interruption, which can significantly affect the thermal history, and consequently, the microstructural features (e.g., grain size) and residual stresses in the vicinity of the interruption plane. In addition, the purge gas flow typically stops after a machine stoppage. Lack of shielding gas could allow oxygen to reach the melt tracks, which could be detrimental to the quality of the part. Particularly, for the LPBF Al-Si-10Mg, the formation of oxide films during processing can be a cause of porosity. Furthermore, improper powder application upon resuming the build can lead to inadequate melting, increasing the chance of lack- of-fusion (LOF) defects.

Furthermore, improper powder application upon resuming the build can lead to inadequate melting, increasing the chance of lack- of-fusion (LOF) defects. This study aims to investigate the effects of various conditions of build interruption (i.e., duration of the interruption and air exposure), surface finish, and heat treatment on the microstructural and mechanical properties of LPBF Al-Si-10Mg and Ti-6Al-4V alloys.2. Experimental ProcedureThe Al-Si-10Mg and Ti-6Al-4V samples for this study were fabricated via an LPBF system (Renishaw AM 400) in a vertical direction. The chemical composition of the utilized powders, provided by the manufacturer (LPW, US), is given in Table 1. Optimized process parameters, recommended by Renishaw, were used for the fabrication of the Al-Si-10Mg and Ti-6Al-4V samples in both interrupted and non-interrupted (i.e., control) conditions. The utilized process parameters are reported in Table 2.

The utilized process parameters are reported in Table 2. Samples were fabricated in dogbone and cylindrical shape to study the effect of surface conditions on the monotonic tensile behavior, as shown in Figures 1(a) and (b), respectively. After the fabrication, cylindrical samples were machined into the dogbone shape, with the same geometry as those directly fabricated in dogbone shape in accordance with ASTM E8/E8M [16]. Dimensions and configurations of the sample, as well as the location of interruption, are illustrated in Figure 1.Table 1. Chemical composition of the Al-Si-10Mg and Ti-6Al-4V powders.MaterialChemical composition (wt. %)Al-Si-10Mg Ti-6Al-4VSi 11.0, Mg 0.35, Fe 0.11, O 0.04, Al balance Al 6.40, V 4.10, C 0.01, H 0.0029, O 0.11, N 0.02, Fe 0.21, Ti balanceTable 2.

Three interruption protocols were implemented to simulate real-life scenarios that might occur during LPBF processing: (i) pausing the machine to refill unplanned powder runout (15 MIN), (ii) an unexpected power outage requiring significant intervention (12 HR), and (iii) a machine that faults for a prolonged period during unattended hours (24 HR) and each protocol was repeated with and without an air exposure during the interruption period [15]. For the Ti-6Al-4V samples, the process interruption lasted for 24 hours, and the machine was paused halfway through the height of the specimens with no air exposure. Tables 3 and 4 present an overview of the interruption procedures for Al-Si-10Mg and Ti-6Al-4V samples.To simulate the interruption, each build was paused at the same layer number of the level of interruption (i.e., at the height of 40mm for Al-Si-10Mg and 45mm for Ti-6Al-4V samples) to place the interruption level within the gage section.

Machine operation was resumed after finishing the interruption protocols (time period and air exposure) for Al-Si-10Mg and Ti-6Al-4V samples, except for 12 HR Al-Si-10Mg samples. In 12 HR samples, since the heated build platform (BP) was turned off during the interruption. Because of the cooling and contraction within the elevator and samples, it was indispensable to adjust the elevator's height to ensure powder spreading was occurring properly. Several test wiper spreads were done to ensure a thin layer height appropriate; spread of powder was evident on each specimen. After this was deemed sufficient, the machine was restarted [14].595Table 3.

After this was deemed sufficient, the machine was restarted [14].595Table 3. Interruption condition for Al-Si-10Mg samples.Experiment groupg M 0 1 - i S - lA1 2 3 4 5 6 7Interruption time Control 12 HR 24 HR 15 MIN 15 MIN 12 HR 24 HRAir exposure No No No Yes No Yes YesHeated BP On Off On On On Off OnSurface finish As-built As-built As-built As-built As-built As-built As-builtHTNo No No No No No NoGroup IDC-A0-HBP 12-A0 24-A0-HBP 15-A1-HBP 15-A0-HBP 12-A1 24-A1-HBPTable 4.

Interruption condition for Ti-6Al-4V samples.Experiment groupV 4 - lA 6 - iT1 2 3 4 5 6 7 8Interruption time Control Control Control Control 24 HR 24 HR 24 HR 24 HRAir exposure No No No No No No No NoHeated BP On On On On On On On OnSurface finish As-built Machined As-built Machined As-built Machined As-built MachinedHTNo No Yes Yes No No Yes YesGroup IDABC MC ABC_HT MC_HT ABI MI ABI_HT MI_HTIn order to study the interaction of heat treatment (HT) with the process interruption effects on the microstructure, residual stress, and consequently the mechanical performance, the Ti-6Al- 4V samples were annealed at 760 °C for 1 hour in an inert Argon atmosphere, followed by air cooling. This procedure allows for stress relief and decomposition of the martensitic phase [17,18].

This procedure allows for stress relief and decomposition of the martensitic phase [17,18]. No heat treatment was done for the Al-Si-10Mg specimens.In order to quantitatively assess the formation of any possible process-induced defects near the interruption location, the gage section of the tensile specimens was scanned using an X-ray computed tomography (XCT) system. Microhardness distribution was measured along the build direction (within the gage section) via a Vickers microhardness tester (Leco Vicker LM-300AT) under a 500-g load for a dwell time of 10 s using a diamond indenter. Monotonic tensile tests were conducted on three specimens in each group under strain-controlled conditions at a 0.001/s nominal strain rate for both Al-Si-10Mg and Ti-6Al-4V. Finally, microstructures and fracture surfaces were observed using scanning electron microscopy (SEM).3. Results and Discussion3.1.

Finally, microstructures and fracture surfaces were observed using scanning electron microscopy (SEM).3. Results and Discussion3.1. Microstructural characterizationThe process interruption may increase the probability of occurrence of process-induced defects, i.e., gas pores and LOFs, as a result of change in the melt pool dynamics and unsustainable bonding between layers. Analysis of our XCT results for Al-Si-10Mg samples showed no meaningful correlation between the formation of defects within the interruption zone and the interruption conditions (i.e., air exposure and interruption time). However, XCT results for Ti-6Al- 4V samples revealed a significant increase in the total volume of defects near the interruption zone.596In order to investigate the changes in microstructural features near the interruption zone, samples were cut ~5 mm above and below the interruption interface.

For Al-Si-10Mg, SEM image shown in Figure 2 displays the microstructure of the interrupted specimen near the interruption location in 24 HR/air exposure condition. A visible line of discoloration can be seen near the interruption plane, which can be attributed to the changes in thermal history due to the process interruption. This feature was observed for all 24 HR specimens. For the Ti-6Al-4V samples, microstructural features in the vicinity of the interruption zone showed no qualitative difference compared to the control counterparts.Figure 2. SEM image of microstructures for Al-Si-10Mg interrupted sample in 24 HR/air exposure condition.Figure 3 shows the SEM micrographs for Ti-6Al-4V samples fabricated in the dogbone and cylinder shapes, before (NHT) and after heat treatment (HT). As can be seen in Figures 3(a) and (b), the characteristic of melt-pool banding patterns was visible only for the non-heat-treated samples.

As can be seen in Figures 3(a) and (b), the characteristic of melt-pool banding patterns was visible only for the non-heat-treated samples. The microstructure of dogbone and cylindrical samples after heat treatment are shown in Figures 3(c) and (d), respectively. The utilized annealing resulted in homogenization of microstructure and removed the interfacial regions.Figure 3. SEM images of microstructures for Ti-6Al-4V interrupted specimens in (a) dogbone_NHT, (b) cylinder_NHT, (c) dogbone_HT, and (d) cylinder_HT conditions.597To explore any unexpected changes in the distribution of microhardness values near the interruption location, several microhardness measurements were taken along the build direction, passing across the interruption level. Results for both Al-Si-10Mg and Ti-6Al-4V samples showed no distinctive changes in the distribution of hardness values near the location of interruption.

Results for both Al-Si-10Mg and Ti-6Al-4V samples showed no distinctive changes in the distribution of hardness values near the location of interruption. The average Vickers microhardness values for the Al-Si-10Mg and Ti-6Al-4V samples were 135 HV and 425 HV, respectively.3.2. Monotonic tensile testMonotonic tensile tests were performed to investigate the potential relationship between the process interruption and tensile properties, including yield stress, ultimate tensile strength, and elongation to failure, as well as the location of failure. The results of the tensile tests for Al-Si-Mg and Ti-6Al-4V specimens are shown in Figures 4 and 5, respectively.For the Al-Si-10Mg specimens, results revealed that yield stress, ultimate tensile strength, and elongation to failure for all the specimens were almost similar, regardless of the interruption time and air exposure, as can be seen in Figures 4(a) and (b).

Values for ultimate tensile strength as well as the location of failure for Al-Si-10Mg samples agree with the results reported by Binder et al. [10].Figure 4. (a) Yield stress, ultimate tensile strength, and (b) elongation to failure values obtained for Al-Si-10Mg specimens.For the Ti-6Al-4V specimens, as seen in Figure 5, the utilized annealing caused an increase in ductility and a decrease in the ultimate tensile strength, where all of the non-heat-treated specimens had higher elongation to failure and lower ultimate tensile strength compared to their heat-treated counterparts. In addition, as can be seen in Figure 5(a), the machined specimens appeared to have higher ultimate tensile strength than the as-built samples, regardless of being affected by the process interruption. As shown in Figure 5(b), the elongation to failure for all interrupted specimens was lower than their control counterparts.

As shown in Figure 5(b), the elongation to failure for all interrupted specimens was lower than their control counterparts. The least elongation to failure belonged to the interrupted specimens with as-built surface condition. The fracture location of one the as-built specimen was coincident with the interrupted plane, as shown in Figure 7.598Figure 5. (a) Yield stress, ultimate tensile strength, and (b) elongation to failure values obtained for Ti-6Al-4V specimens.3.3. FractographySEM micrograph of tensile fracture surfaces of Al-Si-10Mg and Ti-6Al-4V specimens are presented in Figures 6 and 7. For Al-Si-10Mg, almost half of the interrupted specimens, particularly 24 HR air exposure ones, failed near the interruption location. Tensile fracture surfaces and the location of failure for the 24 HR air exposure group are shown in Figure 6.

Tensile fracture surfaces and the location of failure for the 24 HR air exposure group are shown in Figure 6. It can be seen in Figure 6(a) that the location of failure for two out of three specimens in this group is coincident with the interruption level. A large number of gas-pores (shown by arrow) were observed on the fracture surface of the interrupted Al-Si-10Mg specimens that failed at the interruption level, as can be seen in Figure 6(b). However, it cannot be concluded that these gas pores are formed due to the process interruption since other interrupted specimens that did not fail at the interruption plane have the same amount of gas pores on their fracture surface, as can be seen in Figure 6(c).Figure 6.

(a) Location of failure for the Al-Si-10Mg specimens in the 24 HR with air exposure group, along with the SEM images of tensile fracture surface for the specimens failed (b) near and (c) above the interruption plane.599For Ti-6Al-4V specimens, the location of failure for the non-heat-treated specimens in the as-built surface condition was coincident with the location of interruption. This indicates that post- fabrication heat treatment and surface finishing can minimize any potential influence of build interruption on the failure location. The fracture surface of this specimen is presented in Figure 7, where LOF defects and gas pores can be observed on the fracture surface.Figure 7. Tensile fracture surface of non-heat-treated interrupted Ti-6Al-4V specimen in the as- built condition.4.

Tensile fracture surface of non-heat-treated interrupted Ti-6Al-4V specimen in the as- built condition.4. SummaryIn this study, the effects of various conditions of build interruption (i.e., duration of the interruption and air exposure) as well as post-fabrication heat treatment and surface finishing on the microstructural and mechanical properties of LPBF Al-Si-10Mg and Ti-6Al-4V alloys were investigated, and the results were compared with those from a continuous fabrication route. For the Al-Si-10Mg specimens, monotonic tensile tests showed half of the interrupted specimens failed near the interruption location, particularly the specimens with prolonged interruption time and air exposure (24 HR). Tensile and microhardness tests did not show any distinctive changes in the tensile strength and microhardness values as compared to their control counterparts (i.e., non- interrupted).

Tensile and microhardness tests did not show any distinctive changes in the tensile strength and microhardness values as compared to their control counterparts (i.e., non- interrupted). For the Ti-6Al-4V specimens, the tensile location of failure for one of the non-heat- treated specimens in the as-built surface condition (i.e., dogbone specimens) was coincident with the interruption location. In addition, all interrupted Ti-6Al-4V specimens showed a lower elongation to failure as compared to the control specimens. Similar to the Al-Si-10Mg samples, Ti-6Al-4V samples showed no distinctive changes in the distribution of hardness values near the interruption locations. More studies are still required to evaluate the effect of build interruption on other mechanical properties, such as fatigue and creep.6005. References[1] Yadollahi A, Shamsaei N. Additive manufacturing of fatigue resistant materials: Challengesand opportunities. Int J Fatigue 2017;98:14–31.[2] ASTM.

Materials Science and Engineering: A 2018;732:41–52. [6] Malỳ M, Höller C, Skalon M, Meier B, Koutnỳ D, Pichler R, et al. Effect of Process Parameters and High-Temperature Preheating on Residual Stress and Relative Density of Ti6Al4V Processed by Selective Laser Melting. Materials 2019, Vol 12, Page 930 2019;12:930.[5][7] Yadollahi A, Mahtabi MJ, Khalili A, Doude HR, Newman JC. Fatigue life prediction of additively manufactured material: Effects of surface roughness, defect size, and shape. Fatigue Fract Eng Mater Struct 2018;41:1602–14.[8] Marola S, Manfredi D, Fiore G, Poletti MG, Lombardi M, Fino P, et al. A comparison of Selective Laser Melting with bulk rapid solidification of AlSi10Mg alloy. J Alloys Compd 2018;742:271–9. Terrazas-Najera CA, Mayoral FL, Garcia OF, Hossain MS, Espalin D, Fernandez A, et al. Effects of process interruptions on microstructure and mechanical properties of three face centered cubic alloys processed by laser powder bed fusion.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEFFECTS OF PROCESS PARAMETERS AND HEAT TREATMENT ON THE MICROSTRUCTURE AND MECHANICAL PROPERTIES OF SELECTIVE LASER MELTED INCONEL 718Wenpu Huang, Zemin Wang, Jingjing Yang, Huihui Yang, and Xiaoyan ZengWuhan National Laboratory for Optoelectronics Huazhong University of Science and Technology, Wuhan, PR ChinaAbstractIn this study, Inconel 718 superalloy was fabricated by selective laser melting (SLM) and solution treated at 980-1230 ℃ subsequently. The process window was firstly set up based on the density of the samples. Samples were fabricated using various parameters within the process window to investigate the effects of process parameters on microstructure and mechanical properties. The average dendrite arm spacing and the volume fraction of Laves phase raise along with the increasing energy input.

The average dendrite arm spacing and the volume fraction of Laves phase raise along with the increasing energy input. However, no distinct difference of tensile properties was found under parameters in the process window. Interdendritic Laves phase decreases with the solution temperature, while the grain size has the opposite trend. Finally, the solution temperature was fixed at 1080 ℃ to dissolve Laves phases and obtain fine grains. After solution + aging heat treatment, the tensile strengths and ductility all exceed the wrought Inconel 718.Keywords: Selective laser melting; Process parameters; Heat treatment; Microstructure; Mechanical propertiesIntroductionSelective laser melting (SLM) is a rapidly developing advanced manufacturing technique which melts the metal powders with the laser beam to build parts layer by layer directly from CAD model [1, 2].

Compared with conventional manufacturing processes, SLM offers its unique advantages, such as high dimension accuracy and geometrical complexity, material saving and reduction of product cycles [3, 4]. Therefore, the technique has gained widespread interest recently.Inconel 718 (IN718) superalloy is widely applied in aerospace and energy industries due to its excellent strength at elevated temperatures, good corrosion resistance and favorable weldability [5- 7]. Extensive work has been conducted on selective laser melted IN718 at present [1, 2, 4, 8-15]. The microstructure, mechanical properties, heat treatment and residual stress of SLMed IN718 has been studied. However, most of research adopted the optimal process parameters to produce the samples aimed at the minimisation of porosity, and only limited work has focused on the correlation between process parameters, microstructure and properties [11].

Therefore, in this work, effects of process parameters on microstructure and mechanical properties of SLMed IN718 have been studied and the effects of heat treatment were also evaluated.Materials and Methods1268Gas atomized IN718 powders with an average diameter of 31 μm were chosen as the starting materials in this work and the chemical composition is shown in Table 1. All the samples were fabricated by a self-developed SLM machine equipped with an IPG YLR-500 fiber laser (maximum output power of 500 W) .Table1 Chemical composition of IN718 powder Nb Mo 3.04 5.1ElementNiCrFeTi 0.9Al 0.66wt.% Balance 18.46 18.45Co Mn 0.16C0.17 0.054Two kinds of samples were produced: cubic samples (10×10×8 mm3) and tensile test samples. Cubic samples were fabricated using various process parameters shown in Table 2 to optimize the parameters based on the relative density of the samples.

Cubic samples were fabricated using various process parameters shown in Table 2 to optimize the parameters based on the relative density of the samples. Tensile test samples designed according to ISO 6892-1-2009 were built horizontally using the optimal parameters.Table 2 SLM process parameters applied in this study.Process parameter Laser power P (W) Scanning velocity v (mm/s) Hatch spacing h (mm) Layer thickness t (μm)Value 150-500 400-1400 0.08, 0.10, 0.12 20, 40, 60In order to investigate the effects of heat treatment on microstructure and mechanical properties, the as-fabricated samples were solution treated at different temperatures (980-1230 ℃) for 1 h followed by air cooling and conventional double aging treatment (720 °C/8 h + 620 °C/8 h, AC) was adopted. All the heat treatments were carried out using a KSL-1400X-A2 muffle furnace with a heating rate of 10 ºC/min.The relative density of the samples were measured using the Image-Pro Plus 6.0 software.

For microstructure research, the samples were etched by a solvent consisting of 30 ml HCl + 10 ml HNO3 after grinding and mechanical polishing. The microstructure was characterized using a Nikon Epiphot-300 optical microscope (OM), a FEI Sirion 200 scanning electron microscopy (SEM) and a FEI Quanta 200 scanning electron microscopy. Tensile test were conducted at room temperature using a Shimadzu AG-100 kN tester.Results and discussionsEffects of process parametersA series of cubic samples were fabricated using the process parameters shown in Table 2 and the relative density was measured. Based on the results, certain process parameters which guarantee the samples with relative density > 99.9% were selected for further study as shown in Table 3. Sample P1, P2 and P3 only differ in layer thickness and hence they were compared to reveal the effects of layer thickness.

Sample P1, P2 and P3 only differ in layer thickness and hence they were compared to reveal the effects of layer thickness. Similarly, Sample P3, P7 and P9 were compared to reveal the effects of laser power, Sample P4, P5 and P6 were compared to reveal the effects of scanning velocity, Sample P6, P7 and P8 were compared to reveal the effects of hatch spacing.1269Table 3 The selected process parameters which guarantee the samples with relative density > 99.9%.SampleP1 P2 P3 P4 P5 P6 P7 P8 P9Layer thickness t (μm)20 40 60 60 60 60 60 60 60Laser power P (W)300 300 300 400 400 400 400 400 500Scanning velocity v (mm/s) 1000 1000 1000 600 800 1000 1000 1000 1000Hatch spacing h (mm)0.10 0.10 0.10 0.08 0.08 0.08 0.10 0.12 0.10Fig. 1. Microstructures of samples fabricated with different process parameters : (a) P1; (b) P2; (c) P3; (d) P4; (e) P5; (f) P6 (g) P7; (h) P8; (i) P9.1270Fig.

1. Microstructures of samples fabricated with different process parameters : (a) P1; (b) P2; (c) P3; (d) P4; (e) P5; (f) P6 (g) P7; (h) P8; (i) P9.1270Fig. 1 presents the microstructures of IN718 samples produced with different process parameters. It shows that due to the high cooling rates during SLM process, the as-fabricated samples are characterized with fine cellular dendrites with Laves phase precipitating in the interdendritic regions. The formation of Laves phase results from the segregation of alloy elements such as Nb and Mo [16]. In addition, the presence of the phase will weaken the precipitation of γ′ and γ″ [15]. In order to investigate the effects of process parameters on the microstructure, the average dendrite arm spacing and the volume fraction of Laves phase were measured. The binary processing of SEM images were conducted to facilitate the measurements of the volume fraction as shown in Fig. 2.

The binary processing of SEM images were conducted to facilitate the measurements of the volume fraction as shown in Fig. 2. The black regions are γ matrix, while the white regions are Laves phase. The statistic results are shown in Fig.3. As can be seen, the average dendrite arm spacing and the volume fraction of Laves phase raise along with increasing laser power and decreasing scanning velocity and hatch spacing, while the effect of hatch spacing on the average dendrite arm spacing is not distinct. The increasing energy input results in lower cooling rate, which leads to the increase of the average dendrite arm spacing. In addition, Nb elements have more time to diffuse due to the lower cooing rate [13]. As a result, it will contribute to the formation of Laves phase.Fig. 2. Sketch of binary image processingFig. 3.

As a result, it will contribute to the formation of Laves phase.Fig. 2. Sketch of binary image processingFig. 3. Statistic results of (a) the average dendrite arm spacing and (b) the volume fraction of Laves phaseTensile samples were produced using the parameters shown in Table 3 to investigate the effects of process parameters. Results of tensile test are shown in Table 4 and the relationship between mechanical properties and the laser energy density is illustrated in Fig. 4. It can be seen from Fig. 4 that the UTS and YS of different samples do not vary much, while the elongation is volatile and the change law is not apparent.1271The precipitation of the γ′ and γ″ phases is the key to improve the strength of IN718. Although the variation of process parameters can change the microstructure and the cooling rate to a certain extent, the precipitation of strengthening phase is still inhibited [10].

Although the variation of process parameters can change the microstructure and the cooling rate to a certain extent, the precipitation of strengthening phase is still inhibited [10]. The mechanical properties mainly relates to internal defects in such a situation. Therefore, when the relative density is high enough, the mechanical properties are unable to be significantly improved by optimizing the parameters.Table 4 Mechanical properties of samples fabricated with different process parameters YS (MPa) UTS (MPa) 1120 ± 11 847 ± 15 1127 ± 9 833 ± 19 1120 ± 3 832 ± 29 1103 ± 14 778 ± 19 1098 ± 8 799 ± 13 1082 ± 9 773 ± 5 1107 ± 10 815 ± 14 1104 ± 7 816 ± 2 1069 ± 2 760 ± 3Sample P1 P2 P3 P4 P5 P6 P7 P8 P9EL (%) 30.9 ± 0.9 28.9 ± 0.4 23.9 ± 2.5 24.4 ± 2.5 30.2 ± 0.5 34.5± 0.8 28.7 ± 1.2 26.2 ± 1.7 31.3 ± 1.0Fig. 4.

4. Mechanical properties of as-fabricated samples in function of the energy densityEffects of heat treatmentSince no distinct difference of tensile properties was found under various parameters, heat treatments were conducted subsequently. Microstructures of the as-fabricated sample solution treated at different temperatures are shown in Fig. 5. When solution treated below 1080 ℃, the dendrites disappear in some regions as shown in Fig. 5a and b, which indicates that partial recrystallization occurs. When solution temperature rises to 1130 ℃, complete recrystallization has taken place and the grains become coarse and equiaxed.Fig. 6 shows the microstructures of heat-treated samples at high magnification. It can be seen that the microstructures become more homogeneous with temperature increasing. There are still1272some residual particles of the Laves phase in the interdendritic regions when solution treated at 980 and 1030 ℃ as shown in Fig.

There are still1272some residual particles of the Laves phase in the interdendritic regions when solution treated at 980 and 1030 ℃ as shown in Fig. 6a and b, but the Laves phases almost dissolve at higher temperatures.Fig. 5. Microstructure of as-fabricated samples solution treated at different temperatures for 1 h, followed by air cooling. (a) 980 ℃. (b) 1030 ℃. (c) 1080 ℃. (d) 1130 ℃. (e) 1180 ℃. (f) 1230 ℃.1273Fig. 6. SEM images of as-fabricated samples solution treated at different temperatures for 1 h, followed by air cooling at high magnification. (a) 980 ℃. (b) 1030 ℃. (c) 1080 ℃. (d) 1130 ℃. (e) 1180 ℃. (f) 1230 ℃.According to the above results, the solution temperature was fixed at 1080 ℃ to dissolve Laves phases and obtain fine grains. Aging treatment was carried out following the solution treatment. Tensile properties of heat-treated samples are shown in Table 5 and compared with wrought material [15].

Aging treatment was carried out following the solution treatment. Tensile properties of heat-treated samples are shown in Table 5 and compared with wrought material [15]. The results show that an obvious increase of the tensile strength occurs with a decrease in the ductility after heat treatment, but the tensile strengths and ductility all exceed the wrought IN718. Therefore, heat treatment is essential to obtain excellent mechanic properties.1274Table 5 Mechanical properties of as-fabricated and heat-treated samplesSample SLM SLM + heat treatment WroughtYS (MPa) 847 ± 15 1369 ± 17 1034UTS (MPa) 1120 ± 11 1529 ± 19 1276EL (%) 30.9 ± 0.9 18.6 ± 0.9 12ConclusionsIn this study, effects of process parameters and heat treatment on microstructure and mechanicalproperties of selective laser melted Inconel 718 have been studied.

(1) The as-fabricated samples are characterized with fine cellular dendrites with Laves phase precipitating in the interdendritic regions.The average dendrite arm spacing and the volume fraction of Laves phase raise along with the increasing energy input.(2) When the relative density is high enough, the mechanical properties are unable to besignificantly improved by optimizing the parameters.(3) Laves phase decreases with the solution temperature, while the grain size has the opposite trend. The solution temperature was fixed at 1080 ℃ to dissolve Laves phases and obtain fine grains. After solution + aging heat treatment, the tensile strengths and ductility all exceed the wrought Inconel 718.AcknowledgementsThis work is supported by the National Program on Key Basic Research Project of China underGrant no. 613281.References[1] N. Nadammal, S. Cabeza, T. Mishurova, T. Thiede, A. Kromm, C. Seyfert, L. Farahbod, C. Haberland, J.A. Schneider, P.D.

[16] W.M. Tucho, P. Cuvillier, A. Sjolyst-Kverneland, V. Hansen, Microstructure and hardness studies of Inconel 718 manufactured by selective laser melting before and after solution heat treatment, Materials Science and Engineering: A 689 (2017) 220-232.1276.

fects of Process Variables and Size Scale on Solidification Microstructure in Laser-Based Solid Freeform Fabrication of Ti-6Al-4VN.W. Klingbeil, S. Bontha, C.J. Brown and D.R. Gaddam Department of Mechanical and Materials Engineering Wright State University Dayton, OH 45435P.A. Kobryn Materials and Manufacturing Directorate (AFRL/MLSC) Air Force Research Laboratory WPAFB, OH 45433H.L. Fraser Department of Materials Science and Engineering The Ohio State University Columbus, OH 43210J.W. Sears Additive Manufacturing Laboratory South Dakota School of Mines and Technology Rapid City, SD 57701Abstract:This paper summarizes a combination of analytical and numerical modeling approaches which have been used to investigate the effects of process variables and size scale on solidification microstructure in laser-deposited Ti-6Al-4V.

The analytical approach is based on the well-known Rosenthal solution for a moving point heat source, which provides dimensionless process maps for solidification cooling rate and thermal gradient (the key parameters controlling microstructure) as a function of laser deposition process variables (laser power and velocity). Based on these process maps, results for both 2-D thin-wall and bulky 3-D geometries are plotted on solidification maps for predicting grain morphology in laser-deposited Ti-6Al-4V. Although the Rosenthal results neglect the nonlinear effects of temperature- dependent material properties and latent heat of transformation, a comparison with 2-D and 3-D nonlinear thermal finite element results for both small-scale (LENSTM) and large-scale (high power) processes suggests that they can provide reasonable estimates of trends in grain morphology.

Finally, 3-D cellular automaton solidification modeling is used to provide direct predictions of solidification microstructure, and results are compared to experimental observations for both LENSTM and a larger scale process under development at SDSM&T. The results of this work suggest that changes in process variables could potentially result in a grading of the microstructure throughout the depth of the deposit, and that a transition from columnar to equiaxed microstructure is possible at high laser powers.IntroductionThis work is part of a broader effort to tackle the primary obstacles to the widespread commercialization of laser-based additive manufacturing processes, which include the control of melt pool size, residual stress and microstructure [1].

The current study focuses on the prediction of trends in solidification microstructure, with specific application to Ti-6Al-4V.92Laser-based additive manufacturing of Ti-6Al-4V is currently under consideration for application to aerospace components, for which consistent microstructure and resulting mechanical properties are of vital importance [2,3]. However, only limited experimental data exists to link deposition process variables (e.g., laser power and velocity) to resulting microstructure (e.g., grain size and morphology) [4-8], and suitable microstructures have typically been obtained only by trial and error.

In addition, it is unclear whether knowledge based on small-scale laser deposition processes (e.g., LENSTM) can be applied to large-scale (higher power) processes currently under development for commercial application.The current study employs a combination of analytical and numerical modeling approaches to investigate the effects of process variables and size scale on solidification microstructure in laser-deposited Ti-6Al-4V. The analytical approach is based on the well- known Rosenthal solution for a moving point heat source traversing an infinite substrate [9], which has been used recently in the literature to guide the development of process maps relating laser deposition process variables to melt pool size and residual stress [10-13].

Based on the Rosenthal solution, thermal process maps for solidification cooling rate and thermal gradient (the key parameters controlling microstructure) have been presented in [14], and are used herein to provide general insights into the roles of process variables and size scale on microstructure in laser deposited materials.While the Rosenthal solution assumes temperature-independent material response, the nonlinear effects of temperature-dependent properties and latent heat are further included through thermal finite element modeling. Continuum finite element modeling of laser deposition processes is well established, and has been successfully used to study the effects of process variables on melt pool size and residual stress [10-13]. In the current study, the primary purpose of the finite element modeling is to assess the validity of the Rosenthal results for predicting trends in solidification microstructure.

In the current study, the primary purpose of the finite element modeling is to assess the validity of the Rosenthal results for predicting trends in solidification microstructure. This has been investigated for small-scale (LENSTM) deposition of thin-wall geometries in [14-15], and is extended herein to include 3-D finite element modeling of bulky 3-D geometries for both small-scale (LENSTM) and large-scale (higher power) processes. Numerical results extracted from both the Rosenthal and FEM solutions are plotted on solidification maps for predicting grain morphology in Ti-6Al-4V, and the utility of the Rosenthal solution is discussed.Finally, 3-D cellular automaton modeling is used to provide direct predictions of solidification microstructure in laser-deposited Ti-6Al-4V.

Cellular automaton modeling originally developed for casting processes [16] has been recently applied to laser deposition of both titanium and aluminum alloys [15,17], and shows strong potential for linking process variables with microstructure in laser materials processing. In this study, results obtained for a thin-wall geometry are compared to experimental observations for the LENSTM process, while results for a bulky 3-D geometry are compared to experimental observations for a somewhat larger scale process under development at the South Dakota School of Mines and Technology (SDSM&T).The results of this work suggest that size scale can potentially have a significant effect on microstructure, including a transition from columnar to equiaxed microstructure at high laser powers.

These results are particularly relevant in light of recent investigations into the effects of process scaling on the control of melt pool size in laser-based manufacturing processes [18,19].93Geometries ConsideredThe modeling described herein considers both the thin-wall (2-D) and bulky (3-D) geometries of Figure 1, in which the process variables of interest are the absorbed laser power (cid:1)Q and velocity V. In each case, it is assumed that the height h and length L are sufficiently large such that the steady-state Rosenthal solution for a point heat source traversing an infinite half-space applies [9]. In the current study, the geometry of Fig. 1a is used to investigate small- scale deposition of thin-wall structures, such as those commonly manufactured using the LENSTM process. The geometry of Fig.

1a is used to investigate small- scale deposition of thin-wall structures, such as those commonly manufactured using the LENSTM process. The geometry of Fig. 1b is used to investigate both small-scale and large-scale (high power) deposition of bulky 3-D deposits, and provides significant insight into the effect of process size scale on solidification microstructure.αQαQVVxo yoxozozohhLLbba)b)Figure 1. a) Thin-Wall (2-D) and b) Bulky (3-D) Geometries ConsideredRosenthal Solution and Thermal Process MapsThermal process maps for solidification cooling rate and thermal gradient (the key parameters controlling microstructure) have been developed based on the Rosenthal solution for a moving point heat source traversing an infinite substrate [9].

(cid:8)The results are plotted in Figure 2 as a function of normalized melting temperature T m and relative depth within the melt pool z o /z m. The normalized melting temperature varies with laser power, and is defined asT m =Tm (cid:2) To (cid:3)Q.(cid:1)kbIn the above equations, Tm is the melting temperature of the material, To is the initial temperature of the wall, b is the wall thickness, and (cid:2), c and k are the density, specific heat and thermal conductivity of the material, respectively. The normalized depth varies in the range 0 (cid:1) z o /z m (cid:1) 1, where z m signifies the deepest extent of the melt pool for a given value of T m .94(1)(2)The results of Fig. 2 indicate that for a fixed laser velocity, changes in laser power (orchanges in T m ) can have a significant effect on the solidification cooling rate and thermal gradient, and hence the resulting microstructure.

For a given laser power and velocity, the cooling rate varies significantly throughout the depth of the melt pool, while the thermal gradient is insensitive to depth. Moreover, increasing laser power (or decreasing T m ) results in a substantial decrease in thermal gradient throughout the depth of the melt pool, while the cooling rate is most significantly affected at the surface. Hence, increasing laser power (i.e., increasing process size scale) acts to decrease the high thermal gradients typically associated with a columnar microstructure, with an increase in solidification rates toward the surface of the deposit. This suggests the potential for a grading of the microstructure throughout the depth of the deposit, with a transition from columnar to equiaxed microstructure at the surface.∂ T ∂ t=⎛ ⎜⎜ ⎝2 π 2 k b α ρ Q cV2⎞ ⎟⎟ ⎠∂ T ∂ t∇ = T⎛ ⎜⎜ ⎝2π 2 k b α ρ Q cV⎞ ⎟⎟ ∇ T ⎠z 0 zmT m=T m α Q−T oπkbz 0 zmT ma)b)Figure 2.

Process Maps for Solidification a) Cooling Rate and b) Thermal Gradient for Thin-Wall Geometries [14]Changes in laser velocity further affect cooling rate and thermal gradient through the normalizations of eq. (1), which reveal that cooling rate is the more sensitive of the two. However, both the thermal gradient and solidification velocity (the ratio of cooling rate to thermal gradient) scale linearly with laser velocity, which depending on the material system can also influence trends in grain morphology. That said, the laser velocities used in small-scale and large-scale processes can be comparable (on the order of 10-20 mm/s), while the range of laser powers can span as much as two orders of magnitude (300-30,000 W). As such, the remaining discussions on the effects of process scaling will focus primarily on laser power.Process Maps for Bulky 3-D GeometriesAs outlined in [14], the dimensionless solidification cooling rate and thermal gradient for the geometry of Fig.

(cid:7) (cid:10) (cid:8)It should be noted that the above definition of T m is in keeping with that used for bulky 3-D geometries in [14], and differs by a factor of 2 from the definition used by other researchers [18, 20]. It should finally be noted that the presence of the laser velocity V in the temperature normalization of eq. (4) results in a higher order velocity dependence of the cooling rates and thermal gradients of eq. (3), as compared to the 2-D normalizations of eq. (1).∂ T ∂ t=⎛ ⎜ ⎝2 k ρ cV⎞ ⎟ ⎠2⎛ ⎜ ⎝π k α QV⎞ ⎟ ⎠∂ T ∂ t∇ = T⎛ ⎜ ⎝2 k ρ cV⎞ ⎟ ⎠2⎛ ⎜ ⎝π k α Q⎞ ⎟ ∇ T ⎠z 0 zmT m=(αQ− T 0 ρ cVT m) ⎛⎜ ⎝π k2k⎞ ⎟ ⎠z 0 zmT m=a)b)Figure 3. Process Maps for Solidification a) Cooling Rate and b) Thermal Gradient for Bulky (3-D) Geometries [14]The results of Fig.

Process Maps for Solidification a) Cooling Rate and b) Thermal Gradient for Bulky (3-D) Geometries [14]The results of Fig. 3 reveal that for fixed laser velocity, trends in dimensionless cooling rate and thermal gradient are similar to those observed for thin-wall geometries in Fig. 2, except that the dimensionless thermal gradient is slightly more sensitive to depth within the deposit. However, owing to the different temperature normalizations in 2-D and 3-D, comparing the magnitudes of the dimensionless results in Figs. 2 and 3 can be misleading. For example, although the magnitudes of the dimensionless cooling rates are larger in Fig. 2a than in Fig. 3a, the actual cooling rate for a given laser power and velocity is greater in 3-D than in 2-D. In addition, variations in laser velocity for bulky 3-D geometries affect the value of T m , which complicates the interpretation of the axes in Fig. 3.

In addition, variations in laser velocity for bulky 3-D geometries affect the value of T m , which complicates the interpretation of the axes in Fig. 3. However, if it is assumed that the velocity is held constant, all previous conclusions regarding the effect of laser power (i.e., process size scale) on grain morphology are applicable to both 2-D thin-wall and bulky 3-D deposits.96(αQ(3)(4)− T 0 ρ cVT m) ⎛⎜ ⎝π k2k⎞ ⎟ ⎠Thermal Finite Element ModelingWhile the Rosenthal solution assumes temperature-independent material response, the nonlinear effects of temperature-dependent properties and latent heat have been included through thermal finite element modeling. The modeling procedures for the 2-D thin-wall geometry of Fig. 1a are outlined in [15], and have been extended herein to include the bulky 3-D geometry of Fig. 1b. A representative finite element mesh and boundary conditions for a half-symmetric model of a bulky 3-D geometry is shown in Figure 4.

1b. A representative finite element mesh and boundary conditions for a half-symmetric model of a bulky 3-D geometry is shown in Figure 4. The model uses 8-noded bi-linear thermal elements, and has been generated using the software package ABAQUS. The element model finite laser the approximates deposition process as a moving point heat source (cid:1)Q, which is successively applied to adjacent nodes at time intervals corresponding to the laser velocity V. The parameter (cid:1) represents the fraction of the laser power absorbed by the deposit, and has been estimated as 35%. This value gives reasonable agreement with melt pool sizes observed for the LENSTM process [21]. The remaining boundary conditions are insulated (q = 0) on the top and all side faces, and a fixed temperature condition on the bottom (T = 25 ºC). Finally, the finite element model uses temperature-dependent specific heat, density and thermal conductivity, and includes latent heat effects for Ti-6Al-4V.Figure 4.

Finally, the finite element model uses temperature-dependent specific heat, density and thermal conductivity, and includes latent heat effects for Ti-6Al-4V.Figure 4. Representative Thermal Finite Element Mesh for the Bulky 3-D Geometry of Figure 1b.Meshes similar to that of Fig. 4 have been used to investigate both small-scale and large- scale processes, with laser powers ranging from 300-30,000 W . In all cases, the mesh dimensions have been scaled to be in keeping with the steady-state Rosenthal solution, so that the behavior in the vicinity of the melt pool is unaffected by the boundaries. In general, the mesh resolution of Fig. 4 has provided more than 10 elements through the depth of the melt pool, from which solidification cooling rates and thermal gradients have been extracted. Procedures for extracting solidification cooling rates and thermal gradients from the finite element results followed those outlined in [15], and are not reiterated here.

Procedures for extracting solidification cooling rates and thermal gradients from the finite element results followed those outlined in [15], and are not reiterated here. A rigorous convergence study for the case of temperature-independent properties has recovered solidification cooling rates and thermal gradients within one percent of the Rosenthal results of Fig. 3, which verifies the accuracy of the FEM modeling procedures.Solidification Maps for Ti-6Al-4VAs discussed in [14,15], results for solidification thermal gradient and cooling rate can be interpreted in the context of a solidification map to provide predictions of grain morphology in97laser-deposited Ti-6Al-4V. Given the solidification cooling rate (cid:1)T /(cid:1)t and thermal gradient G = (cid:1)T , the solidification velocity R is determined asR =1 G(cid:1)T (cid:1)t.The expected grain morphology can be predicted as either equiaxed, columnar or mixed by plotting points in G vs.

R space (i.e., on the “solidification map”), which has been previously calibrated for Ti-6Al-4V [7].Solidification maps showing the effect of laser power for small-scale (LENSTM) deposition of bulky 3-D geometries are shown in Figure 5. Analogous results for thin-wall geometries are presented in [14]. The results of Fig. 5a are extracted directly from the 3-D Rosenthal results of Fig. 3, with thermophysical properties for Ti-6Al-4V assumed constant at the melting temperature Tm=1654°C. The results of Fig. 5b are extracted from the thermal finite element model of Fig. 4, with temperature-dependent properties and latent heat effects for Ti- 6Al-4V. The range of powers is 350-750 W, which is larger than that considered for thin-wall geometries in [14].

4, with temperature-dependent properties and latent heat effects for Ti- 6Al-4V. The range of powers is 350-750 W, which is larger than that considered for thin-wall geometries in [14]. In each case, the laser velocity is held constant at V = 8.47 mm/s, and the fraction of absorbed laser power is taken to be (cid:1)=0.35.Small-Scale (LENSTM): 3-D Rosenthal (V = 8.47 mm/s)Small-Scale (LENSTM): 3-D FEM ( V = 8.47 mm/s)100000100000Q = 350 W Q = 550 W Q = 750 WIncreasing Z (Depth)Q = 350 W Q = 550 W Q = 750 W1000010000Fully ColumnarFully Columnar1000Mixed1000MixedG (K/cm)Fully EquiaxedG (K/cm)Fully Equiaxed10010010101 0.0010.010.1 R (cm/s) a)1101 0.0010.010.1R (cm/s) b)1Figure 5.

Predicted Ti-6Al-4V Grain Morphology for Small-Scale (LENSTM) Deposition of Bulky 3-D Deposits from a) 3-D Rosenthal Solution and b) 3-D FEMAlthough the Rosenthal results neglect the nonlinear effects of temperature-dependent properties and latent heat of transformation, trends in G vs. R predictions are in good agreement with those obtained from the FEM results. In particular, both the Rosenthal and FEM results predict a fully columnar morphology, which is in keeping with experimental observations of LENSTM deposited Ti-6Al-4V [5-7]. However, results also suggest that increasing the laser power tends to shift the data closer to the boundary for a mixed columnar/equiaxed grain98(5)10morphology. These results are in keeping with G vs.

These results are in keeping with G vs. R predictions reported for thin-wall geometries in [14], as well as with experimental observations recently reported in [4].A comparison of solidification map predictions from the Rosenthal and FEM solutions for large-scale (high power) processes is shown in Figure 6. The range of laser powers considered spans 5-30 kW. In each case, the laser velocity is held constant at V = 8.47 mm/s, and the fraction of absorbed laser power is taken to be (cid:1)=0.35. As observed for small-scale processes, trends in G vs. R predictions from the 3-D Rosenthal and FEM solutions are in good agreement. In particular, both suggest that large-scale processes can result in a grading of the microstructure throughout the depth of the deposit, with a mixed or even fully-equiaxed microstructure at the surface.

Moreover, the trend toward equiaxed microstructure increases with laser power.Large-Scale Processes: 3-D Rosenthal (V = 8.47 mm/s)Large-Scale Processes: 3-D FEM (V = 8.47 mm/s)10000010000010000Q = 5000 W Q = 10000 W Q = 15000 W Q = 20000 W Q = 30000 WIncreasing Z (Depth)10000Q = 5000 W Q = 10000 W Q = 15000 W Q = 20000 W Q = 30000 W1000Fully Columnar1000Fully ColumnarG (K/cm)100MixedFully EquiaxedG (K/cm)100MixedFully Equiaxed10101 0.0010.010.1 R (cm/s) a)1101 0.0010.010.1 R (cm/s) b)1Figure 6. Predicted Grain Morphology for Large-Scale (High-Power) Deposition of Bulky 3-D Deposits from a) 3-D Rosenthal Solution and b) 3-D FEMThe above trends can be inferred from the previous discussion of the thermal process maps of Figs. 2 and 3. In particular, increasing laser power (i.e., size scale) acts to reduce the thermal gradients, which for a fixed cooling rate would move the data down and to the right in G vs. R space.

2 and 3. In particular, increasing laser power (i.e., size scale) acts to reduce the thermal gradients, which for a fixed cooling rate would move the data down and to the right in G vs. R space. However, increasing power also decreases cooling rates (and hence the solidification rate R), which is a competing effect. The net result is an essentially downward movement in G vs. R space. At the same time, cooling rates increase toward the surface of the deposit, and are accompanied by a slight decrease in thermal gradients for bulky 3-D geometries (Fig. 3). This also tends to move the data down and toward the right in G vs. R space.

3). This also tends to move the data down and toward the right in G vs. R space. The net result is a trend toward mixed or equiaxed microstructure at the surface, which increases with laser power (i.e., process size scale).9910Cellular Automaton Modeling of Solidification MicrostructureCellular automaton (CA) modeling has been used to provide direct predictions of solidification microstructure in laser-deposited Ti-6Al-4V, for both 2-D thin-wall and bulky 3-D geometries.

Results for 2-D thin-wall geometries are compared to experimental observations the LENSTM process, while results for a bulky 3-D geometry are compared to experimental observations for a slightly larger-scale process under development at SDSM&T.fora)b)The general procedures used here follow those outlined in [15], and involve 3-D thermal finite element modeling of laser deposition process using the software package ProCastTM, followed by 3-D cellular automaton solidification modeling of grain nucleation and growth using the software package CAFE3D. Detailed background on the algorithms used in CAFE3D can be found in [16]. In brief, the software uses a Gaussian distribution of nucleation sites to simulate the stochastic grain growth process during solidification, and requires statistical solidification parameters for the material system of interest.

These parameters, including the mean undercooling, the standard deviation of the undercooling, and the nucleation site density, have been calibrated by direct comparison with bulk solidification of cast Ti-6Al-4V ingots of varying size [21].thec)d)e)f)Figure 7. CA Predicted and Observed Microstructures in Thin-Wall LENSTM Deposits (Q=370 W, V=12.5 mm/s)Modeling of Thin-Wall LENSTM DepositsModeling procedures and results for a thin-wall LENSTM geometry are shown in Figure 7. For a given set of process variables (laser power and velocity), thermal history output from a ProCastTM model of the laser deposition process (Fig 7a) is used as input for subsequent CAFE3D cellular automaton analysis of grain nucleation and growth. However, since the CAFE3D software is designed to model solidification from an entirely liquid state, it is necessary100a)c)Laser Directionb)d)face analyzedto first isolate the melt pool region from the ProCastTM model.

As illustrated in Fig. 7b, thermal history data from the section of the model within the melt pool (T > 1650 ºC ) is extracted from the ProCastTM model, and used to define an equivalent melt pool model suitable for CAFE3D analysis. initial temperatures and boundary conditions for the melt pool model of Fig. 7b result in largely one-dimensional heat transfer through the depth of the melt pool, which approximates the directional solidification occurring in laser deposition processes.The500 WattsThe evolution of the solidification microstructure is shown in Figs. 7c -d, and a comparison of CAFE3D predictions and observed the microstructures LENSTM process (Q=370 W, V=12.5 m m / s ) is shown in Figs. 7e-f. The grain size and morphology of the CAFE3D predictions are in reasonable agreement with the observed microstructure.

7e-f. The grain size and morphology of the CAFE3D predictions are in reasonable agreement with the observed microstructure. Both are fully columnar, which is in keeping with solidification map predictions based on both the Rosenthal and 2-D FEM solutions [14,15], as well as previous observations for LENSTM deposition of Ti-6Al-4V [5-7].500 Watts 500 Watts300 (cid:1)m 300 (cid:1)m 300 (cid:1)mfor300 _me)f)g)h)Figure 8. Predicted Microstructure in Bulky 3-D Deposits and Comparison with Laser Glazing at SDSM&T (Q=500 W, V=20 mm/s)Modeling of Bulky 3-D GeometriesSimilar procedures have been used to model solidification microstructure for bulky 3-D geometries, as shown in Figure 8. Thermal history data from the half-model of Fig. 8a is used to define the equivalent melt pool model of Fig. 8b.

Thermal history data from the half-model of Fig. 8a is used to define the equivalent melt pool model of Fig. 8b. Compared to the melt pool model for the 2-D thin-wall geometry, the only difference for the bulky 3-D geometry is the lack of convection and radiation conditions associated with the wall faces. These conditions are replaced by thermal insulation, which results in truly one-dimensional heat conduction through the depth of the melt pool. Evolution of the resulting microstructure for a laser power of 500 W is shown in Figs. 8c-101d. The result is fully columnar, which is in keeping with the solidification map predictions from both the Rosenthal and 3-D FEM results.Model results are compared to experimental observations for laser glazing experiments conducted at SDSM&T in Figs. 8e-h. This is a somewhat larger-scale process than LENSTM, and uses a CO2 laser with powers on the order of 500-1750 W. The results of Fig. 8 are for a laser power of 500 W and velocity of 20 mm/s.

8e-h. This is a somewhat larger-scale process than LENSTM, and uses a CO2 laser with powers on the order of 500-1750 W. The results of Fig. 8 are for a laser power of 500 W and velocity of 20 mm/s. The red melt pool region of Fig. 8e assumes an absorbed laser power of (cid:1)=0.35, which gives reasonable agreement with the observed melt pool dimensions of Fig. 8f. A comparison of the predicted and observed microstructures is shown in Figs. 8g-h. Both microstructures are columnar, which is in keeping with the solidification map predictions previously discussed.To date, the CAFE3D modeling has been successful in predicting the fully columnar microstructures associated with laser deposition at relatively small scales, including both the LENSTM and SDSM&T processes. Similar modeling at higher laser powers is underway, and seeks to confirm the columnar to equiaxed transition predicted by the solidification map approach.

Similar modeling at higher laser powers is underway, and seeks to confirm the columnar to equiaxed transition predicted by the solidification map approach. It should be noted that there are recent experimental results in the literature to support such a transition, which has been reported for a 14 kW large-scale process in [5].Summary and ConclusionsThis work has employed a variety of analytical and numerical modeling approaches to investigate the effects of process variables and size scale on solidification microstructure in Ti- 6Al-4V. Numerical results based on the Rosenthal solution for a moving point heat source have been plotted on solidification maps for predicting grain morphology in laser-deposited Ti-6Al- 4V, and are in keeping with results from nonlinear thermal FEM analyses for both small-scale and large-scale (high power) processes.

Results suggest that size scale can have a significant effect on grain morphology, with a transition from columnar to mixed/equiaxed microstructure at high powers. Finally, cellular automaton modeling has been used to provide direct predictions of solidification microstructure in both thin-wall and bulky 3-D deposits, and results are in keeping with experimental observations for both the LENSTM and SDSM&T processes.AcknowledgementsThis work has been supported by the National Science Foundation, grant number DMI-0224517, by the Joint AFRL/DAGSI Research Program, project number ML-WSU-01-11, and by the U.S. Army Research Laboratory and U.S. Army Office under grant number DAAD 19-02-2-001.References 1. Beuth, J.L. and Klingbeil, N.W., 2001, “The Role of Process Variables in Laser-Based Direct Metal SolidFreeform Fabrication,” JOM, Vol. 53, No. 9, pp. 36-39.2. Kobryn, P.A. and Semiatin, S.L., 2001, “The Laser Additive Manufacture of Ti-6Al-4V,” JOM, Vol. 53, No. 9, pp.

Birnbaum, A., Aggarangsi, P. and Beuth, J., 2003, “Process Scaling and Transient Melt Pool Size Control in Laser-Based Additive Manufacturing Processes,” Solid Freeform Fabrication Proceedings, (D.L. Bourell, R. H. Crawford, J.J. Beaman, K.L. Wood, H.L. Marcus, eds.), Austin, August 2003.20. Aggarangsi, P., Beuth, J.L., and Gill, D.D., 2004, “Transient Changes in Melt Pool Size in Laser Additive Manufacturing Processes,” Solid Freeform Fabrication Proceedings, Austin, August 2004 (in the current proceedings).21. Brown, C.J., 2003, “Modeling of Solidification Microstructure in Laser-Deposited Ti-6Al-4V,” Masters Thesis, Department of Mechanical & Materials Engineering, Wright State University, June 2003.103.

fects of Processing Conditions on Prototypes Reinforced with TLCPs for Fused Deposition ModelingR. W. Gray N, D. G. Baird, and 1. H. BahnDepartment ofChemical Engineering,Department ofMechanical Engineering,Rapid Prototyping Laboratory, andPolymer Materials and Interfaces LaboratoryVirginia Polytechnic Institute and State UniversityPolypropylene (PP) composite strands, reinforced with thermotropic liquid crystalline polymers (TLCPs), were generated using a novel dual extrusion process which allowed for the use of a TLCP with a significantly higher melting temperature than that of the PP. Pregenerated TLCPtpp microcomposite strands were reprocessed using a second novel process to produce a wen-controned monofilament composite for use in a FDM 1600 rapid prototyping system in order to build complex geometries.

Uniaxial parts were built to determine the effect of differing material compositions and processing temperatures, in order to develop an oPerating window for the optimal mechanical properties. By adjusting the lay down pattern of orientable materials, the final mechanical properties of the part could be engineered independent ofthe material. To understand the effect of the reprocessing steps on the pregenerated microcomposites, the final mechanical properties ofthe monofilament composite were compared with those ofthe pregenerated strands.1.0 IntroductionAt present, only a few polymeric materials with limited mechanical properties are used in fused deposition modeling (FDM) rapid prototyping systems. Many of the prototypes fabricated can only serve as geometric replicas ofthe proposed production part because ofthe inherently poor mechanical properties.

Many of the prototypes fabricated can only serve as geometric replicas ofthe proposed production part because ofthe inherently poor mechanical properties. The materials that are commercially available for the FDM 1600 rapid prototyping system are acrylonitrile butadiene styrene copolymer (ABS), a nylon copolymer, and investment casting wax. Of the three commercially available materials, ABS has the highest tensile modulus and strength, and the properties of prototypes fabricated using ABS have been determined to be 1.5 GPa and 22 MPa, respectively [1]. Therefore, there is interest in developing materials that can be used to fabricate prototypes with higher mechanical properties which give the parts greater functionality.Thermotropic liquid crystalline polymers (TLCPs) are a novel class of materials that have potential for use in FDM applications for several reasons.

First, it has been shown that TLCPs have excenent tensile properties with moduli ranging from 50 GPa to It has also been shown that due to their excenent tensile 100 GPa for neat fibers.449properties and due to their fibril forming nature, TLCPs have been used to reinforce thermoplastics [2 - 4]. The diameter of the reinforcing TLCP fibrils are typically one order of magnitude smaller compared to typical glass and carbon fiber. However, it may be not possible to extrude glass and carbon fibers through the die head and still maintain high aspect ratio fiber (i. e. un > 100) due to the small diameter capillary die used in order to fabricate dimensionally precise prototypes.The final mechanical properties of TLCP based composites, where the reinforcing fibrilar morphology is developed during the processing step on an in situ basis, are directly dependent upon the processing conditions [4 - 7].

Shear and extensional flow fields during processing serve to deform dispersed TLCP droplets into reinforcing fibrils and impart molecular orientation to the fibril, which results in increased tensile properties in the direction of the flow field. This preferred orientation, of both the molecules and the fibrils, results in anisotropic mechanical properties. It is generally known that extensional flow fields develop higher aspect ratio fibrils and greater molecular orientation, and hence extensional flow yields higher tensile properties than strong shear flow. For example, 20 wt% Vectra A, a commercial TLCP marketed and sold by Hoechst Celanese, was used to reinforce polypropylene (PP), and composites that were generated via fiber spinning and injection molding had tensile moduli of 9.6 GPa and 2.6 GPa, respectively [5, 8, 9]. Therefore, in order to obtain the optimal reinforcement, strong extensional forces, such as those present in fiber spinning, are necessary.

Therefore, in order to obtain the optimal reinforcement, strong extensional forces, such as those present in fiber spinning, are necessary. The purpose of this work is to fabricate TLCP reinforced composite prototypes in such a fashion as to capitalize on the full potential ofthe reinforcement.2.0 Experimental 1.1 MaterialsThe TLCP used was Vectra A950 (Hoechst Celanese). It is a random copolyester based on hydrobenzoic acid (73% mole) and 2-hydroxy-6-napthoic acid (27% mole). It has a glass transition temperature of 108°C and a melting point of 283°C [10]. However, Vectra A needs to be heated to 320°C to melt out all of the residual crystallites [11]. The matrix material used was the 4018 grade of Amoco polypropylene (PP).

However, Vectra A needs to be heated to 320°C to melt out all of the residual crystallites [11]. The matrix material used was the 4018 grade of Amoco polypropylene (PP). It has a melting point of 160°C and a melt flow index (MF) of 13.5 [12].2.2 Spinning of TLCPIPP Composite StrandsA novel dual extrusion process was used to blend the TLCP and PP forming a self reinforced composite fiber in order to obtain optimal mechanical properties in the the thermoplastic and TLCP are composite [13 - 15]. plasticated in separate extruders so that independent thermal histories are imposed to the two materials. TLCP is heated to sufficiently high temperatures to fully melt all of the crystallites. For example, the melting point ofVectra A is 283°C, but it needs to be heated to 3200C in order to fully melt all the residual crystallites [11].

For example, the melting point ofVectra A is 283°C, but it needs to be heated to 3200C in order to fully melt all the residual crystallites [11]. Vectra A is then supercooled to the temperature at which the matrix is processed, and using a multiple-port injection nozzle is then injected into the matrix stream which results in continuous TLCPIn this extrusion process,450streams encapsulated within the matrix. The melt is passed through a series of static mixers which serve to divide the TLCP streams, before being extruded through a capillary die and drawn in the spinline to high draw ratios in order to achieve high levels of molecular orientation in the TLCP fibrils.There are several advantages of the dual extrusion process over generating TLCP fibrils in situ during FDM. Processing the TLCP in a separate extruder minimizes degradation of the matrix because the matrix is not exposed to the high temperatures that are necessary to fully melt the TLCP.

Processing the TLCP in a separate extruder minimizes degradation of the matrix because the matrix is not exposed to the high temperatures that are necessary to fully melt the TLCP. The TLCP is supercooled before being injected into the matrix minimizing the degradation of the matrix. The dual extrusion process does not rely on droplet deformation because the multiple-port injection nozzle introduces continuous TLCP streams into the matrix, resulting in high aspect ratio fibrils. Finally, the composite strands can be post-processed above the melting point of the matrix and below the melting point ofthe TLCP using a variety of conventional processing techniques (e. g.

Finally, the composite strands can be post-processed above the melting point of the matrix and below the melting point ofthe TLCP using a variety of conventional processing techniques (e. g. FDM) which allow for the retention of the TLCP reinforcing fibrils generated during the dual extrusion process in the final part [1, 8, 9, 14 - 23].3.0 Discussion and Results 3.1 Novel Process for Generation of Composite Feed Stock for FDMIn order to develop new materials for use in the FDM 1600 rapid prototyping system that could be used to fabricate prototypes with greater mechanical properties and self-reinforced thermoplastic greater composite strands were generated using a novel dual extrusion process. Vectra A, a TLCP know for its exceptional mechanical properties, was used as the reinforcing phase, and PP served as the matrix.

Vectra A, a TLCP know for its exceptional mechanical properties, was used as the reinforcing phase, and PP served as the matrix. Vectra AlPP strands were spun with 20 wt% and 40 wtOlO TLCP concentrations, and the moduli of the strands were 9.55 GPa and 22.8 GPa, Further results concerning the properties of these strands are given respectively. elsewhere [8, 9].functionality than those presently available,Strands from the dual extrusion process were then post-processed to form monofilaments with a well controlled diameter for use in the FDM 1600 rapid prototyping In this process, the continuous TLCP reinforcement was granulated into short system. TLCP fibrils having lengths less than 6 mm and diameters ranging from 1 to 5 f.lm. The tensile moduli of the 20 wtOlO and 40 wt% monofilaments were 1.9 GPa and 2.2 GPa, respectively, and the strengths were 25.6 MPa and 21.1 MPa, respectively.

The tensile moduli of the 20 wtOlO and 40 wt% monofilaments were 1.9 GPa and 2.2 GPa, respectively, and the strengths were 25.6 MPa and 21.1 MPa, respectively. The properties compare favorable to those of pure PP where the tensile modulus and strength are 0.98 GPa and 23.2 MPa, respectively [15]. The Vectra A reinforcement resulted in an increase in modulus of approximately 1000,4 over those neat PP, while having similar strengths. The tensile properties of the monofilaments were significantly lower than those of the strands from the dual extrusion process. Previous work has shown that these lower properties were probably due to a reduction in aspect ratio, poor fibril distribution, and poor fibril alignment [20 - 22].4513.2 Plaque Fabrication via FDMVectra AlPP monofilaments were used as feed stock to fabricate plaques via FDM, and the effects of fabrication temperature on the tensile properties of the final part were examined. Monofilaments with 20 wtOA. and 40 wtOA.

Monofilaments with 20 wtOA. and 40 wtOA. Vectra A were used to build parts with a lay-down pattern aligned uniaxially in the machine direction. In Figure 1, the tensile modulus is shown as a function offabrication temperature for both the 20 wtOA. and 40 wtOA. Vectra A composite parts. Plaques with 20 wtOA. Vectra A reinforcement were fabricated at processing temperatures of 190°C, 240°C, and 290°C. The moduli of these plaques were essentially independent of processing temperatures and were approximately 1.6 GPa, even above the melting point of the reinforcement. The 40 wtO~ Vectra A monofilaments were processed at 240°C and 290°C, and the moduli of the plaques were 2.7 GPa and 2.4 GPa, respectively. Thus, the moduli of the fabricated plaques decreased somewhat when processed above the melting point of the reinforcement.

Thus, the moduli of the fabricated plaques decreased somewhat when processed above the melting point of the reinforcement. Parts were not fabricated at 190°C because the FDM 1600 rapid prototyping system was unable to generate the pressure required to extrude the composite.In Figure 2, the tensile strength is shown as a function of fabrication temperature. The strengths of both the 20 wt% and 40 wt% Vectra A reinforced plaques decreased by approximately one third when processed at 290°C as compared to the strengths of the plaques when processed below the melting point ofVectra A The strength of the 20 wt% Vectra A plaques decreased from approximately 33 MPa when fabricated below the melting point of the Vectra A to 21 MPa when fabricated above the melting point. Similarly, the strength of the 40 wt% Vectra A plaques decreased from 37 MPa when post-processed at 240°C to 25 MPa when extruded above the melting point of the TLCP.

Similarly, the strength of the 40 wt% Vectra A plaques decreased from 37 MPa when post-processed at 240°C to 25 MPa when extruded above the melting point of the TLCP. When processed below the melting point of Vectra A, the composite relies on the reinforcement generated in the dual extrusion process. However, when Vectra AlPP composites are processed above the melting point of the TLCP, the reinforcing fibrils are melted and ofientation is lost within the TLCP phase. In the melt state, the fibrils form into droplets, and there are minimal extensional forces present in the FDM 1600 rapid prototyping system to develop the fibrilar morphology required for the optimal tensile properties.

Thus, tensile properties of VectraIPP prototypes, where reinforcement was generated in the dual extrusion process, were somewhat better than those where the reinforcement was generated during prototype fabrication.The mechanical properties of the Vectra AlPP monofilament feed stock were compared to those of the plaques fabricated via the FDM 1600 rapid prototyping system. The modulus of the 40 wt% Vectra A monofilament was measured to be 2.2 GPa. Plaques fabricated uniaxially in the machine direction at 240°C had a modulus of 2.7 GPa. the strength of the composite increased from 21 MPa to 37 MPa in the Similarly, fabricated plaque. The modulus of the 20 wtOA. Vectra A monofilaments and plaques was similar before and after fabrication and was approximately 1.8 GPa. The strength of the 20 wt% Vectra A monofilaments increased from 26 MPa to 33 MPa after being fabricated into a plaque.

The strength of the 20 wt% Vectra A monofilaments increased from 26 MPa to 33 MPa after being fabricated into a plaque. The increase in mechanical properties of the 40 wt% VeetTa A composite was probably due to a change in alignment of the reinforcing fibrils as a result of the flow452It has been shown elsewhere that the tensile kinematics of the FDM 1600 extruder. properties of post-processed long fiber composite strands are very sensitive to the flow It was shown that a critical un was necessary in order to kinematics of the die [8, 9]. obtain the optimal mechanical properties and that the extrusion rate, and die diameter can effect the tensile properties ofthe extrudate. Vectra A1PP (28/72 wtOA») composite strands from the dual extrusion process were post-processed into monofilaments, and the maximum tensile modulus and strength were 4.0 GPa and 45 MPa.

Vectra A1PP (28/72 wtOA») composite strands from the dual extrusion process were post-processed into monofilaments, and the maximum tensile modulus and strength were 4.0 GPa and 45 MPa. Thus, higher properties can be obtained with lower Vectra A concentrations under the proper post processing conditions [8, 9].Plaques were fabricated from composite monofilaments using three different lay down patterns: uniaxial in the machine direction, uniaxial in the transverse direction, and a 0-90 lay-down pattern. The plaques were fabricated from the Vectra A1PP (40/60 wtOA») composite monofilaments at 240°C which was the processing temperature that resulted in the greatest tensile properties for plaques fabricated uniaxially in the machine direction. In Figure 3, the mechanical properties are shown as a function of lay-down pattern. The abscissa is defined as the volume percent of material laid in the machine direction.

In Figure 3, the mechanical properties are shown as a function of lay-down pattern. The abscissa is defined as the volume percent of material laid in the machine direction. Thus, plaques fabricated entirely in the transverse direction had 0 vol% laid in the machine direction, and parts with 0-90 lay-down patterns had 50 vol% laid in the machine direction, and parts built uniaxially in the machine direction had 100 vol% of the material It is shown that both the tensile modulus and strength laid in the machine direction. increased monotonically with the volume percent laid-down in the machine direction. For parts fabricated uniaxially in the transverse direction and parts fabricated uniaxially in the machine direction, the modulus increased from 1.3 GPa to 2.7 GPa, and the strength increased from 10 MPa to 37 MPa.

This dependence of mechanical properties on the lay down pattern was due to anisotropic reinforcement of the matrix by the TLCP fibrils as was observed in Figure 4, and this reinforcement was aligned in the direction in which the roads were laid. This was why the greatest mechanical properties were observed in the plaques fabricated uniaxially in the machine direction. Plaques fabricated uniaxially in the transverse direction had the lowest properties because the reinforcement was aligned orthogonal to the machine direction. Thus, the bulk properties of these plaques were highly dependent on the tensile properties of the matrix. Also, the mechanical properties of a part built in the transverse direction were probably somewhat lower due to weld lines perpendicular to the machine direction. As predicted by composite theory, there was a monotonic increase in mechanical properties with increasing volume fraction laid-down in the machine direction [24].

As predicted by composite theory, there was a monotonic increase in mechanical properties with increasing volume fraction laid-down in the machine direction [24]. Therefore, using composite theory, the mechanical properties ofthe final part can be engineered to match the requirements ofthe proposed prototype by adjusting the lay-down pattern.4.0 ConclusionsPlaques were fabricated from Vectra A1PP composites and neat Vectra A monofilaments via FDM. The tensile modulus of Vectra A1PP (40/60 wtOA») composite than those of ABS prototypes and plaques were approximately 100% greater453approximately 1500A» greater than those of pure PP, and these properties could probably be increased iflong fiber reinforced composite monofilaments were used as feed stock.in TLCPs and TLCP Due to the anisotropic mechanical properties present composites, the lay-down pattern affected the properties of the final part.

It was found that the tensile properties of Vectra A1PP composites increased monotonically with roads laid in the machine direction. Thus, the final mechanical properties of a prototype can be increasing the tailored to a specific application by adjusting the lay-down pattern, functionality ofthe prototype.5.0 AcknowledgmentsThe authors would like to gratefully recognize the support from the Army Research Office (Grant No. DAAH04-94-G-0282), the Naval Surface Warfare Center, Dahlgren Division. and Armstrong Laboratory, Brooks AFB under contract N60921-89 DA239, Order 0059, and Stratasys, Inc. Also, the authors would like to thank Prof: G. L. Wilkes for the use ofthe SEM and M. A McLeod for operating it.1 2R W. Gray IV, D. G. Baird, I. H. Behn, Rap. Proto. J., in preparation (1997). A. I. Isayev, in Liquid-Crystalline Polymer Systems Technological Advances, ed. A. I. lsayev, T. Kyo, and S. Z. D. Cheng, American Chemical Society, Washington D. C. 1996. D. G.

fects of Processing on Microstructure and Mechanical Properties of Ti-6Al-4V Fabricated using Electron Beam Melting (EBM): Orientation and Location*Nikolas Hrabea, Ryan Kircherb, Timothy QuinnaaNational Institute of Standards and Technology (NIST), Boulder, CO bMedical Modeling Inc., Golden, COOfficial contribution of the National Institute of Standards and Technology; not subject to copyright in the United States.ABSTRACTElectron beam melted (EBM) titanium alloy (Ti-6Al-4V) samples were built and characterized (qualitative prior-β grain size, quantitative α lath thickness, monotonic tensile, Vickers microhardness) to determine the effect of location and orientation on microstructure and properties. Samples of vertical orientation, compared to horizontal, were found to have 30% lower elongation.

Samples of vertical orientation, compared to horizontal, were found to have 30% lower elongation. Orientation within the x-y plane as well as location were found to have less than 3 % effect on mechanical properties, and it is possible a second order effect of thermal mass contributed to these results.INTRODUCTIONProcessing variables for selective electron beam melting (EBM) can be divided into two main categories: inter-build and intra-build. Inter-build variations occur across multiple builds, whereas intra-build variations occur within one build space and potentially within one part. For EBM as well as other additive manufacturing techniques it is important to understand the effects both types of these variables have on microstructure and properties in order to achieve material of acceptable quality, repeatability, and reproducibility.

Examples of inter-build processing variables that have been investigated previously are chemistry (both starting powder and as-built material) [1] and build plate temperature [2].The present work focuses on intra-build variations and specifically location and orientation. For location we hypothesized that relative temperature would decrease radially outward to the exterior of the build space, and it was expected that as relative temperature decreased, cooling rate and time above β transus would increase. For the expected Ti-6Al-4V microstructure (known as acicular, lamellar, or Widmanstätten), as cooling rate decreases, α lath thickness is known to increase, leading to lower ultimate tensile strength (UTS), yield strength (YS), and microhardness [3-9]. Another microstructural feature characterized in this work, prior-β grain size, is controlled by time above the β transus.

Another microstructural feature characterized in this work, prior-β grain size, is controlled by time above the β transus. For the expected Ti-6Al-4V microstructure, as time above the β transus increases, prior-β grain size increases, leading to lower UTS, YS, and microhardness [3-9]. Considering the known microstructure-mechanical property relationships for the expected Ti-6Al-4V microstructure, it was expected that exterior parts would have finer microstructure (both α laths and prior-β grains) and higher mechanical properties (UTS, YS, microhardness) compared to parts on the interior. The part geometry1045developed for this work provides the opportunity to isolate the location variable, and to the knowledge of the authors, this has not been done previously.Orientation was the second intra-build variable investigated, and two separate comparisons were made: horizontally oriented vs. vertically oriented as well as x-oriented vs. y- oriented parts.

vertically oriented as well as x-oriented vs. y- oriented parts. For EBM processed Ti-6Al-4V, elongated prior-β grains and microstructural texture has been previously observed in the z-direction [2], and it was expected that these factors would be the dominant mechanism responsible for any observed effects of orientation. Horizontally oriented and vertically oriented parts have been compared previously, including two studies on EBM processed Ti-6Al-4V. The first [10] reported no difference in UTS or YS, and the horizontal part displayed higher percent elongation at break (% EL). The second reported no difference in YS, but UTS was higher and % EL lower for the horizontal part [11]. There was not enough part layout information provided in the second study to speculate on reasons for the difference in trends.

There was not enough part layout information provided in the second study to speculate on reasons for the difference in trends. Classically annealed bulk Ti-6Al-4V, of the same microstructure as that expected for EBM Ti-6Al-4V, has previously been reported to exhibit increasing % EL with decreasing prior-β grain size [5, 6]. The mechanism reported was reduced slip length for the shorter prior-β grain boundary lengths of smaller prior-β grains lead to shorter dislocation pile-up lengths, reduced stress concentrations at triple-point grain boundary intersections, delayed crack initiation, and increased % EL. Considering these bulk material results and the elongated prior-β grain morphology in the z-direction for EBM Ti-6Al-4V [2], horizontally oriented parts were expected to have higher % EL compared to vertically oriented parts because horizontally oriented parts would have shorter prior-β grain boundary lengths in the tensile loading axis.

In [12] tensile data was collected for Ti-6Al-4V fabricated using electron beam freeform fabrication (EBF3), a process similar to EBM. No intra-build comparison to work presented here can be made because all the specimens were taken from the same part, and the heat transfer dynamics of the wire fed EBF3 differ from that of the powder bed process of EBM.Orientation effects within the x-y plane were also investigated in this work, and as the tensile axes were all orthogonal to the elongated prior-β grain and microstructural texture direction, we expected no difference would be observed in microstructure or mechanical properties.MATERIALS AND METHODSAll parts were built in one batch, and the position and orientation of each part in the 200 mm x 200 mm build space (centered on the 210 mm x 210 mm build plate) was chosen to isolate each intra-batch variable and minimize effects from other variables (Figure 1).

Tensile, microstructure, and microhardness specimens were cut from the as-built parts and prepared for testing. After testing, results were analyzed to determine any variation in microstructure or mechanical properties as a function of energy input, orientation, and location.Arcam S12 EBM equipment1 (software version 3.2.45.12281, accelerating voltage 60 kV, layer thickness 70 µm, “standard” Arcam build theme for Ti-6Al-4V) was used with gas atomized Ti-6Al-4V powder (average particle size 70 µm). This build theme varies electron beam parameters in a controlled sequence throughout the build according to algorithms developed by the manufacturer in an effort to achieve fully dense as-built parts with consistent1046microstructure and properties. Part 1 was built with a speed factor of 30. Parts 2-6 were built with a speed factor of 40. This range of speed factor represents the limits within which complete melting and fully dense as-built parts can be expected.

Parts 2-6 were built with a speed factor of 40. This range of speed factor represents the limits within which complete melting and fully dense as-built parts can be expected. Speed factor is a manufacturer-specific variable that describes the more generic electron beam parameters such as beam current and beam speed as they relate to a given manufacturer-specific build theme. Therefore, specific values of beam current and beam speed cannot be known from the speed factor because they are constantly being varied by the build theme throughout the build. Since speed factor has little practical use for additive manufacturing systems other than the one used in this work, it should be thought of in terms of its inverse relationship with energy input.Adjacent parts had 1 mm space between them, which was assumed to be enough space to thermally isolate each part. The dimensions of horizontally built parts 1, 2, and 4 were the same (x-dimension 159 mm, y-dimension 14 mm, and z-dimension 27 mm).

The dimensions of horizontally built parts 1, 2, and 4 were the same (x-dimension 159 mm, y-dimension 14 mm, and z-dimension 27 mm). Horizontally built part 6 had the same dimensions but a different orientation (x-dimension 14 mm, y-dimension 159 mm, and z-dimension 27 mm). Parts 3 and 5 also had the same dimensions but were built in the vertical orientation (x-dimension 27 mm, y-dimension 14 mm, and z-dimension 159 mm).Figure 1. Top view schematic of part layout/orientation in 200 mm x 200 mm build space. A 1 mm spacing was used between adjacent parts.Chemistry was measured for the as-built parts and compared to ASTM F2924 Standard Specification for Additive Manufacturing Titanium-6 Aluminum-4 Vanadium with Powder Bed Fusion. Aluminum, vanadium, and iron were measured by optical emission spectroscopy (OES). Oxygen, nitrogen, and hydrogen were measured by inert gas fusion. Carbon was measured by the combustion method.

Aluminum, vanadium, and iron were measured by optical emission spectroscopy (OES). Oxygen, nitrogen, and hydrogen were measured by inert gas fusion. Carbon was measured by the combustion method. Approximate uncertainties in each measurement were as follows: oxygen (3 %), nitrogen (14 %), hydrogen (6 %), and carbon (6 %). Uncertainties for aluminum, vanadium, and iron were not available.Bulk tensile samples (Figure 2b) were cut from as-built parts (Figure 2a) by use of electric discharge machining (EDM). Each tensile sample was sliced by use of EDM into ten thin1047tensile specimens (Figure 2c) used for monotonic tensile testing. Results from previous work by this author [13] showed negligible variation of microstructure and properties within a given part as a function of distance from the build plate. Therefore, for each part, mechanical properties will be reported as an average of the values for each of the 10 tensile specimen slices.

Therefore, for each part, mechanical properties will be reported as an average of the values for each of the 10 tensile specimen slices. Tensile specimens were approximately equal to the rectangular cross-section subsize specimens of ASTM E8 Standard Test Methods for Tension Testing of Metallic Materials with the following gauge dimensions: length 25 mm, width 5 mm, and thickness 2 mm (±0.1 mm tolerance for all dimensions). By use of the orientation nomenclature from ASTM 2921 Standard Terminology for Additive Manufacturing – Coordinate Systems and Test Methodology, tensile specimens for each part had the following orientation: parts 1, 2, and 4 – XY; parts 3 and 5 – ZY; part 6 – YX. The remaining material after cutting the bulk tensile sample (Figure 2d) was sectioned for metallographic mounting (Figure 2e), and the same piece was used to characterize microhardness and microstructure.

It is important to note that this microhardness/microstructure piece is directly adjacent to the tensile specimen gauge sections, allowing appropriate comparisons between microstructure, microhardness, and mechanical properties. Nonparametric ANOVA (Kruskal-Wallis) statistical analysis was performed for comparisons between parts or between distances from the build plate within a given part. P-values less than 0.05 were considered significant.Figure 2. (a) as-built part, (b) bulk tensile sample cut from as-built part, (c) tensile specimens after slicing tensile sample, (d) remnant after cutting bulk tensile sample, (e) smaller piece cut from remnant used for microhardness and microstructure characterization. All cutting operations performed by use of EDM.Monotonic tensile testing was performed according to ASTM E8 for all specimens at a strain rate of 10-3 sec-1.

All cutting operations performed by use of EDM.Monotonic tensile testing was performed according to ASTM E8 for all specimens at a strain rate of 10-3 sec-1. Yield strength (YS, 0.2 % offset method), elongation at break (% EL), and ultimate tensile strength (UTS) were measured from the engineering stress- strain curve of each specimen. Tensile specimen fracture surfaces were observed using scanning electron microscopy (SEM). The microstructure/microhardness piece of each part (Figure 2e) was mounted and polished to a 0.02 µm finish after removal of EDM layer. Ten Vickers microhardness measurements (1000gf, 55x, 20 second hold, >2.5 indentation width spacing between adjacent indentations) were taken at distances equivalent to the middle of each of the 10 tensile specimens for a total of 100 indentations per part.

Microstructure/microhardness pieces were then etched (Kroll’s etchant) and both prior-β grain size as well as α lath thickness were1048observed under an optical microscope. For the expected Ti-6Al-4V microstructure, prior-β grain size increases with increasing time above the β transus, leading to lower UTS, YS, and microhardness [3-9]. This microstructural feature was qualitatively assessed by use of lower magnification images taken under low-angle light. This light was necessary to achieve sufficient contrast between grains, but it also resulted in inconsistent lighting, leaving the top of most images darker than the bottom. It is important to note that this gradient of brightness is only due to the low-angle light and should not be mistaken as a microstructural feature. Quantitative measurements were not possible due to lack of grain boundary definition (α did not segregate to all grain boundaries).

Quantitative measurements were not possible due to lack of grain boundary definition (α did not segregate to all grain boundaries). α lath thickness was quantified with the use of Image J1 [14] and higher magnification optical microscope images (Figure 3). Three images were taken adjacent to the microhardness indentations for each represented tensile specimen, and 20 α laths were measured for each image totalling 600 α laths measured for each part. It is important to note that the quantification method chosen depends on the subjective judgment of the person making the measurement, and any potential measurement errors were minimized by having a single person make all measurements.Figure 3. Representative high magnification optical microscope image of EBM Ti-6Al-4V microstructure (light phase = α, dark phase = β), including measurements of α lath thickness (black lines, some denoted by white arrows).

The lines were drawn by the operator, and Image J automatically recorded their length.RESULTSMicrostructureThe microstructure of EBM processed Ti-6AL-4V (Figure 3) observed in this work is similar to other EBM Ti-6Al-4V [11] but differs slightly from the expected acicular or Widmanstätten microstructure obtained through classical annealing [8]. The classical morphology is discrete α in continuous β, but in the EBM microstructure of this work it appears α is continuous and β is discrete.1049ChemistryChemistry is reported to help understand the observed magnitudes of mechanical properties. The measured chemistry of as-built parts (Table I) met ASTM F1472, except for oxygen. Out-of-specification oxygen content in as-built parts is due to out-of-specification oxygen in the powder (0.21 wt %). Failure of any parts in this study to meet mechanical property requirements might be attributed to high oxygen content.

Failure of any parts in this study to meet mechanical property requirements might be attributed to high oxygen content. Out-of-specification oxygen was acceptable for this work because parts within the same build were compared, and if it is assumed chemistry does not change within a build, then any effects of intra-build processing variables could still be observed.Table I. Measured as-built part chemistry (in wt %) was out-of-specification for oxygen compared to additive manufactured Ti-6Al-4V with powder bed fusion. Measurement uncertainties for iron, aluminum, and vanadium were not available.

Measurement uncertainties for iron, aluminum, and vanadium were not available. Out-of-specification oxygen content in as-built parts is from out-of-specification oxygen content in powder (0.21 wt %).NCHFeOAlVASTM F2924as-built0.05 max0.10 max0.015 max0.031±0.004 0.019±0.001 0.0008±0.000050.30 max 0.20 max 5.5-6.75 3.5-4.5 4.070.20.22±0.0076.1Location – horizontalWe hypothesized that the exterior of the build space might be cooler and lead to higher cooling rates than the interior of the build space. Therefore, it was expected that compared to interior part 2, exterior part 4 would exhibit finer α laths and higher UTS and YS. Parts 2 and 4 were compared as they have the same orientation and volume. Also, the compact nature of the part geometry developed for this work allows for excellent isolation of the location variable and an accurate assessment of any effect of this variable on microstructure and mechanical properties.

Although results (Table II) suggest the exterior part did have finer α laths, UTS (1 % change), YS (2 % change), and microhardness were all lower than for the interior part. There was no observed difference in % EL or prior-β grain size (Figure 4).Table II. Properties of horizontal location parts. A large change in location resulted in a small change in UTS (1 % change) and YS (2 % change).partdescriptionUTS (MPa)YS (MPa)% ELα lath thickness (µm)microhardness (HV)interior 2 exterior 4 ANOVA p-value1029.7±7.0 982.9±5.7 12.2±0.8 0.95±0.31 372.0±7.2 1017.4±4.9 966.5±5.3 12.2±2.0 0.91±0.24 360.1±8.1 0.0003<0.0001>0.9999 0.02<0.00011050Figure 4. Macroscopic optical microscope images of (a) interior part 2 and (b) exterior part 4 showing prior-β grain structure. The bottom of each image corresponds to the bottom of the part, adjacent to the build plate. The top of each image corresponds to the top of the part.

The bottom of each image corresponds to the bottom of the part, adjacent to the build plate. The top of each image corresponds to the top of the part. The gradient of brightness is only due to the low-angle light used to capture the image and should not be mistaken as a microstructural feature. It appears there is little difference in grain size.Location – verticalVertically oriented parts isolate the location variable even better than horizontal parts in the x-y plane, so interior part 3 and exterior part 5 were compared to investigate any trends. The same relationship between location and microstructure/mechanical properties was expected, but UTS (2 % change) and YS (2 % change) were again lower for the exterior part compared to the interior part (Table III). No differences in % EL, α lath thickness, microhardness, or prior-β grain size (Figure 5) were observed.

No differences in % EL, α lath thickness, microhardness, or prior-β grain size (Figure 5) were observed. Voids were observed on the fracture surfaces of 7 out of 10 tensile specimens for part 5 (Figure 6), compared to only 2 out of 10 tensile specimens for part 3. These voids were not observed on any of the horizontally built parts. In the x-y plane voids ranged in size (approximate minimum 150 µm) but were larger than the internal pores (approximate maximum 100 µm, Figure 6a,b,c) that originate from hollow gas-atomized powder and remain in as-built material [11]. In the z-direction voids were relatively thin (approximate maximum 20 µm, Figure 6d,e). Voids were also non-spherical compared to spherical internal pores. For all voids, unidirectional lines with the same orientation (left to right in Figure 6a) were observed.

Voids were also non-spherical compared to spherical internal pores. For all voids, unidirectional lines with the same orientation (left to right in Figure 6a) were observed. From observing the microstructure piece for part 5 (Figure 6d,e) voids occurred over a range of distances from the build plate, and seemed, qualitatively to decrease in number from the left side to the right side. There were no voids observed on the microstructure piece of part 3, so the trend observed in part 5 could not be verified.1051Table III. Properties of vertical location parts. A large change in location resulted in a small change in UTS (2 % change) and YS (2 % change).partdescription UTS (MPa)YS (MPa)% ELα lath thickness (µm)microhardness (HV)interior 3 5 exterior ANOVA p-value1032.9±12.9 984.1±8.5 9.0±2.9 0.96±0.26 367.6±8.3 1008.6±15.2 961.0±7.1 7.1±3.4 0.96±0.27 368.4±5.8 0.2 0.0013<0.0001>0.99990.69Figure 5.

Macroscopic optical microscope images of (a) interior part 3 and (b) exterior part 5 showing prior-β grain structure. The left edge shown in each image corresponds to the left edge of each part, and the right shown in each image corresponds to the right edge of each part. The gradient of brightness is only due to the low-angle light used to capture the image and should not be mistaken as a microstructural feature. It appears there is little difference in grain size.1052Figure 6. (a) and (b) optical microscope images (c) SEM image of part 5 tensile specimen fracture surface with 2.6 % EL showing top view of voids (black arrows) and internal pores (white arrows) (d) and (e) optical microscope images of EBM Ti-6Al-4V part 5 microstructure after etching showing side view of voids (black arrows) and internal pores (white arrows). Unidirectional lines were observed in the same orientation on all voids (left to right on (a) and (b)).Orientation – horizontal vs.

Unidirectional lines were observed in the same orientation on all voids (left to right on (a) and (b)).Orientation – horizontal vs. verticalThe z-oriented texture in EBM Ti-6Al-4V [2] suggested microstructure and mechanical properties might change depending on part orientation. Horizontal part 2 and vertical part 3 were compared because the orientation of their tensile axes to texture direction differed. The influence of location was minimized as the parts were adjacent to each other, and both parts were the same volume. % EL was found to be considerably lower (30 % change) for the vertical part compared to the horizontal part (Table IV). Microhardness was also found to be lower for the vertical part despite no observable difference in UTS or YS. There was also no observable difference in α lath thickness or prior-β grain size (Figure 7).

There was also no observable difference in α lath thickness or prior-β grain size (Figure 7). It is also important to point out that the prior-β grain morphology of these parts is elongated in the z-direction as expected from previous EBM Ti-6Al-4V literature [2].1053Table IV. Properties for horizontal vs. vertical orientation parts. Vertical part had 30% lower %EL. No difference was observed for UTS or YS.partdescription UTS (MPa)YS (MPa)% ELα lath thickness (µm)microhardness (HV)horizontal 2 3 vertical ANOVA p-value1029.7±7.0 1032.9±12.9 984.1±8.5 9.0±2.9 0.72 0.5982.9±5.7 12.2±0.8 0.95±0.31 372.0±7.2 0.96±0.26 367.6±8.3 0.650.010.02Figure 7.

Macroscopic optical microscope images of (a) horizontal part 2 (bottom of image corresponds to the bottom of the part, adjacent to the build plate; top of image corresponds to the top of the part) and (b) vertical part 3 (left edge in image corresponds to the left edge of part; right edge in image corresponds to the right edge of part) showing prior-β grain structure. The gradient of brightness is only due to the low-angle light used to capture the image and should not be mistaken as a microstructural feature. It appears there is little difference in grain size.Orientation – x vs. yIt was desirable to compare two parts of different orientation but with the same relationship between tensile axes and texture direction. X-oriented part 4 and y-oriented part 6 meet these criteria and although they are not adjacent, they are both on the edge of the build space.

X-oriented part 4 and y-oriented part 6 meet these criteria and although they are not adjacent, they are both on the edge of the build space. The hypothesized location effect would scale with distance from the center of the build space and be measured radially outward to the exterior of the build space. Therefore, any location effects were still minimized with these two parts as they were both at the exterior of the build space. The y-oriented part has a higher UTS (2 % change), YS (3 % change), and microhardness and a smaller α lath thickness (Table V). There was no observable difference in % EL, but it appeared prior-β grain size was smaller for the y-oriented part (Figure 8).1054Table V. Properties for y vs. x orientation parts.

The gradient of brightness is only due to the low-angle light used to capture the image and should not be mistaken as a microstructural feature. It appears the y-oriented part has a smaller prior-β grain size.DISCUSSIONLocation – horizontalThe magnitude of change in UTS (1 % change) and YS (2 % change) was relatively small despite a large difference in location (Table II). Compared to the interior part, the exterior part had lower UTS, YS, and microhardness as well as a smaller α lath thickness. These relationships do not follow the known relationships between α lath thickness and mechanical properties for the acicular Ti-6Al-4V microstructure [3-9]. One possible explanation is that the relationship between α lath thickness and mechanical properties is different for EBM processed Ti-6AL-4V because its microstructure differs slightly from the expected acicular or Widmanstätten microstructure through classic annealing.

This microstructure difference is explained in more detail in part 1 of this publication. The macroscopic images of prior-β grain size (Figure 4) facilitate only qualitative assessment and may not be an appropriate method for distinguishing small differences in grain size. The mechanical property results are considered more trustworthy and will be used to identify any trends.1055It was hypothesized that UTS and YS would be higher for the exterior part due to lower heat and faster cooling rates from beam deflection. However, the opposite trend was observed (Table II). It is possible that the small change in UTS and YS made it impossible to accurately determine any trends. If it is assumed the trend in the data is accurate, the second order effect of thermal mass might explain the results.

If it is assumed the trend in the data is accurate, the second order effect of thermal mass might explain the results. The 1 mm spacing between adjacent parts was assumed to thermally isolate parts, but it is possible adjacent parts added heat and slowed cooling rates, leading to coarser microstructure and lower UTS and YS than expected. The exterior part (#4, Figure 1) in this case was part of a larger thermal mass (including the two parts “not characterized”) compared to that of the interior part (#2 and #1).Location – verticalSimilar to horizontally built location parts, for vertical location parts a large difference in location resulted in a small change in UTS (2 %) and YS (2 %) (Table III). The trend observed for the vertical location parts was opposite of the expected trend, but it is quite possible the small different in UTS and YS makes accurate identification of any trends impossible.

If the trend is assumed to be accurate, the second order effect of thermal mass does not help in this case because the interior part 3 had a higher thermal mass than the exterior part. The higher frequency of voids found in the exterior part (Figure 6) may have resulted in the observed UTS and YS trends. Out of spec oxygen content (Table I) and/or presence of large voids (Figure 6) could possibly have resulted in % EL for both parts failing to meet the requirement in ASTM F2924 (10 % minimum). If only tensile specimens without voids are considered, % EL for both parts meet ASTM F2924 requirements (10.2 % ± 1.6 % for part 3, 11.1 % ± 0.8 % for part 5). It is well documented for electron beam cold hearth melting (EBCHM) that aluminum vaporizes during melting of Ti-6Al-4V [15, 16]. This occurs for additive manufacturing EBM as well, leading to aluminum condensing and solidifying on parts of the build chamber such as the heat shields.

This occurs for additive manufacturing EBM as well, leading to aluminum condensing and solidifying on parts of the build chamber such as the heat shields. Although the exact mechanism of void formation is unclear, the morphology of the voids leads us to hypothesize that aluminum-rich flakes are flaking off the heat shields and falling into the powder bed where they contribute to the formation of the observed voids.Orientation – horizontal vs. verticalHorizontally built part 2 had a significantly larger % EL (30 % change) compared to the vertically built part 3, and no difference was found for UTS or YS (Table IV). The difference in % EL is attributed to differences in tensile axis orientation with respect to direction of elongated prior-β grains and microstructural texture [2]. These results agree well with classically annealed bulk Ti-6Al-4V [5, 6] and one of the similar previous studies on EBM Ti-6Al-4V [10].

These results agree well with classically annealed bulk Ti-6Al-4V [5, 6] and one of the similar previous studies on EBM Ti-6Al-4V [10]. Unfortunately, there was insufficient methodology and part layout information provided in the other similar previous study [11] to speculate on any reasons for the difference in % EL trend observed.Orientation – y vs. xResults for these parts (Table V) showed no difference in % EL, but there was a smalldifference in UTS (2 % change), YS (3 % change), microhardness, α lath thickness, and prior-β grain size. Orientation was not expected to have an effect in the x-y plane as the tensile axes for parts oriented in both directions were orthogonal to the texture direction [2]. It is possible the second order effect of thermal mass resulted in the observed differences.

It is possible the second order effect of thermal mass resulted in the observed differences. X-oriented part 4 was1056adjacent to larger parts (“not characterized”, Figure 1) which could have slowed the cooling rate, leading to the observed coarser α laths and prior-β grain and lower UTS and YS compared to the more thermally isolated x-oriented part 6.CONCLUSIONSVertically oriented parts were found to have a considerably lower % EL (30 % change) compared to horizontally oriented parts, which was attributed to differences in tensile axis orientation in relation to the direction of elongated prior-β grains and microstructural texture for EBM Ti-6Al-4V, and no difference in UTS or YS was observed. Orientation within the x-y plane and location were both found to have less than 3 % effects on mechanical properties, and a second order effect of thermal mass may have influenced these results.

Orientation within the x-y plane and location were both found to have less than 3 % effects on mechanical properties, and a second order effect of thermal mass may have influenced these results. This study provides a framework for identifying and evaluating intra-build variations using a unique part geometry which allows for more accurate determination of influence from various processing variables. Variables should be evaluated for each new material, machine, and technique (e.g. EBM powder bed different from EBM wire feed) as part of a quality control process.ACKNOWLEDGMENTSThis research was performed while the author held a National Research Council ResearchAssociateship Award at the National Institute of Standards and Technology.FOOTNOTES1Commercial names are identified in order to specify the experimental procedure adequately.

e Effects of Processing Parameters on Defect Regularity in Ti-6Al-4V Parts Fabricated By Selective Laser Melting and Electron Beam MeltingHaijun Gong*, Khalid Rafi*, Thomas Starr†, Brent Stucker*Department of Industrial Engineering, †Department of Chemical Engineering, J. B. Speed School of Engineering, University of Louisville, Louisville, KY 40292 Accepted August 16th 2013AbstractProcessing parameter has an important effect on Selective Laser Melting (SLM) and Electron Beam Melting (EBM) processes. Defects are easily formed by deviating from optimized processing parameters. This study purposely fabricated Ti-6Al-4V specimens with defects by varying process parameters from the factory default settings in both SLM and EBM equipment. Specimen’s density was measured based on the Archimedes method for estimating porosity. Microscopy of specimen’s top surface were observed to compare melt pool and overlap.

Specimen’s density was measured based on the Archimedes method for estimating porosity. Microscopy of specimen’s top surface were observed to compare melt pool and overlap. “Marginal Parameters” is identified to describe the processing parameters which are capable of fabricating specimens with certain porosity. As a result, a correlation between defect regularity and marginal parameters has been established. The effect of marginal parameters on the melt pool is discussed to explain defect formation.1. IntroductionMetal powder bed fusion based Additive Manufacturing (AM) processes, such as Selective Laser Melting (SLM) and Electron Beam Melting (EBM), are successfully being used to manufacture parts from metallic materials including stainless steel, maraging steel, cobalt chromium, titanium alloys and more.

In particular, additively manufactured Ti-6Al-4V parts are of wide interest for aerospace, biomedical and industrial fields due to its fracture resistance, fatigue behavior, corrosion resistance and biocompatibility [1].An SLM machine selectively melts metallic powder layers, forming a solid part on a building plate. Inside an SLM build chamber, there is a material dispenser platform along with a recoating unit used to feed new powder over the build platform, as shown in Fig. 1. Parts are then built up additively, layer by layer, based upon a sliced CAD representation of the desired part.LaserRecoaterRecoater bladeBuilding platePartMetallic PowderZ-axisCollector PlatformBuild PlatformDispenser PlatformFig. 1 Schematic of SLM ProcessIn the EBM process, instead of a laser beam, an electron beam is generated via a tungsten filament in an electron gun, as shown in Fig. 2.

1 Schematic of SLM ProcessIn the EBM process, instead of a laser beam, an electron beam is generated via a tungsten filament in an electron gun, as shown in Fig. 2. The powder bed is first pre-heated using a defocused electron beam in order to enhance the conductivity and cohesion of particles, and to avoid powder spreading in the chamber. The microstructural characteristics of EBM-built material are also influenced by the pre-heating process. A more concentrated beam of electrons is then used to melt the metal powder according to the proscribed CAD data. When the electrons penetrate the powder surface and into the powder grains, their velocity is slowed. By doing this, their kinetic energy is converted into thermal energy and the metal powder reaches its melting temperature [2].

By doing this, their kinetic energy is converted into thermal energy and the metal powder reaches its melting temperature [2]. Electromagnetic lenses are used to control the electron beam, in contrast to the scanning galvanometers used in SLM.Electron gunFocusing lensDeflection coilsPowderPowderRakeBuild specimenBuild tableFig. 2 Schematic of EBM ProcessIn recent years, SLM and EBM of Ti-6Al-4V have made significant progress [3-6], especially related to optimization of processing parameters and characterization of Ti-6Al-4V materials. However, many authors indicated that SLM and EBM-built materials are subject to porosity inclusion, which influences mechanical properties [7, 8]. In order to explain defect generation, this study carried out experiments with various processing parameter combinations. For clarity, a number of non-optimal parameter combinations were named “marginal parameters”.

For clarity, a number of non-optimal parameter combinations were named “marginal parameters”. The effect of marginal parameters on the variation of melt pool and hatch overlap is discussed by comparing the top surface of specimens.2. Ti-6Al-4V powder propertiesIn this study, Raymor Ti-6Al-4V powder was used in the SLM process. Currently, no literature reports on the optimized processing parameters for this specific fine powder, compared to EOS-provided powder. Fine powder is not suitable for the EBM process because small particles will be easily influenced by electrostatic charge, which results in serious powder spreading abovethe building platform [9]. Therefore, Arcam Titanium Ti-6Al-4V powder was used for the EBM process.2.1 Raymor Ti-6Al-4V powder for SLMRaymor Ti-6Al-4V (Grade 23) powder has an apparent density of 2.55 g/cm3. Most particles have diameter no more than 45 µm (0-25 µm, 34.1%; 25-45, µm 63.1%; +45 µm, 2.8%). Fig.

Most particles have diameter no more than 45 µm (0-25 µm, 34.1%; 25-45, µm 63.1%; +45 µm, 2.8%). Fig. 3 shows the powder morphology under a Scanning Electron Microscope (SEM). The powder was measured using a Microtrac S3000 laser-based particle size analyzer. The spherical particles have a size distribution between 17.36 µm (D10) and 44.31 µm (D90) with Mean Volume Diameter around 30 µm. Its particle size is nearly normally distributed around two peaks, with diameter 5.46 µm and 28.67 µm. Raymor Ti-6Al-4V powder contains a lot of fine particles, compared to EOS Titanium Ti-6Al-4V powder [10].Fig. 3 Raymor Ti-6Al-4V Powder Morphology and Particle Size Distribution2.2 Arcam Ti-6Al-4V powder for EBMArcam Titanium Ti-6Al-4V ELI powder has a spherical morphology, with apparent density no more than 2.7 g/cm3. The particle size analysis showed that Arcam powder has a size distribution between 46.94 µm (D10) and 99.17 µm (D90) with Mean Volume Diameter around 72.69 µm.

The particle size analysis showed that Arcam powder has a size distribution between 46.94 µm (D10) and 99.17 µm (D90) with Mean Volume Diameter around 72.69 µm . The particle size is nearly normally distributed. But it has a small percentage of particles with diameter larger than 100 µm, as shown in Fig. 4.Fig. 4 Arcam Ti-6Al-4V Powder Morphology and Particle Size Distribution3. Design of ExperimentsEnergy density is usually used to describe the average applied energy per volume of material during a powder bed fusion process. Significant processing parameters are included in an energy density equation. For example, energy density can be expressed by equation (1) for the SLM process [11].PEvht(1)where P is laser power, v is scanning speed, h is hatch spacing and t is layer thickness. Generally, these parameters have a complicated mutual effect during the laser melting process when all of parameters are varied in an experiment.

Generally, these parameters have a complicated mutual effect during the laser melting process when all of parameters are varied in an experiment. The melt pool variation and defect generation will be easily characterized if some parameters are kept constant. According to the equation (1), it is noted that laser power and scanning speed have an inverse impact on the energy density. So this experiment adopted a factorial design based on these two parameters in order to correlate the energy density and defect regularity. An EOS M270 Direct Metal Laser Sintering (DMLS) system was utilized to fabricate specimens.

An EOS M270 Direct Metal Laser Sintering (DMLS) system was utilized to fabricate specimens. The experimental factors and levels are given in Table 1.Table 1 Factors and Levels of Factorial Experiment (SLM: Ti-6Al-4V)Factor Laser Power (W)Level 40, 80, 120, 160Scanning Speed (mm/s)120, 240, 360, …, 1560Hatch Spacing (µm) Layer Thickness (µm)100 30For the EBM process, “energy density” can also be expressed similarly as equation (1), by substituting electron beam power for laser power [9]. An Arcam S400 EBM system was used for this study. However, the scanning speed cannot be directly varied in the Arcam system. It is dynamically controlled through a “speed function” during the melting process in order to achieve the correct melt pool size. The speed function is related to the “max current” and “speed function index”. Additionally, “line offset” and “focus offset” are also influential processing parameters to the EBM process.

The speed function is related to the “max current” and “speed function index”. Additionally, “line offset” and “focus offset” are also influential processing parameters to the EBM process. Therefore, this study utilized max current, line offset, focus offset, and speed function index as the primary experimental factors. As shown in Table 2, three levels for each factor were selected for studying their significance on the defects through a Taguchi experiment (L9). The One-Factor-at-A-Time (OFAT) method was also employed to further investigate the significant factors.Table 2 Factors and Levels of Taguchi Experiment (EBM: Ti-6Al-4V)FactorMax Current (mA) Line Offset (mm) Focus Offset (mA) Speed Function IndexLevel 10, 20, 30 0.15, 0.20, 0.25 5, 10, 15 60, 120, 180After SLM and EBM specimens were built, each specimen’s density was measured using the Archimedes method [12] according to ASTM B962-08.

The measured density was compared with the nominal density of the Ti-6Al-4V material to estimate porosity according to equation (2),1eDnR(2)where RD is porosity, ρe is measured density, and ρn is nominal density. It can be seen that, the lower the measured density, the larger the porosity.4. Results and DiscussionIn this section, the cause of defect generation during SLM and EBM processing is analyzed and compared. Processing parameters are further classified based on their effect on the porosity generation. Microscopy of specimen’s top surface illustrates the change of melt pool and overlap between hatch lines.4.1 Effect of Processing Parameters on SLM Ti-6Al-4V 4.1.1 Porosity Distribution and ANOVAFig. 5 shows the porosity of SLM produced Ti-6Al-4V specimens with uncertainty value. From Fig. 5(a), it can be seen that there is no apparent porosity when scanning speed is between 600 mm/s and 1600 mm/s at a laser power of 160 W.

From Fig. 5(a), it can be seen that there is no apparent porosity when scanning speed is between 600 mm/s and 1600 mm/s at a laser power of 160 W. When the scanning speed is less than 600 mm/s, which is a higher energy density, the increased porosity implies internal defects in the specimens. Moreover, porosity is not available when the scanning speed is decreased below 360 mm/s, because specimens were not successfully built. The deposited layer was seriously deformed resulting in crash with the recoating blade. When laser power is at 120 W in Fig. 5(b), slow scanning speed (<360 mm/s) also causes build failure. There could be no porosity or internal defects, when the scanning speed ranges from 720 to 1200 mm/s.

5(c). Lowering the scanning speed leads to a small porosity. But the porosity is dramatically increased when increasing the scanning speed. When the laser power is lowered to 40 W, all specimens are assumed to contain defects even if at a slow scanning speed. At 1200 mm/s, the specimen’s porosity is even higher than 20%, as shown in Fig. 5(d).ANOVA of factorial experiment is summarized in Table 3. It can be seen that the F values of both laser power and scanning speed are larger than the threshold (F0.05, 3, 160 and F0.05, 12, 160) which indicate significance with 95% confidence level. This conforms to the results of other references with respect to SLM.SourceDegree of freedomTable 3 ANOVA of Factorial Experiment Adj. Sums of Squares (SS)Sums of Squares (SS)Adj.

6. The energy density of each point in the process window could be simply represented by slope of a line between the origin and a point on the graph, if the hatch spacing and layer thickness are constant. For example, all points along the dash line in Fig. 6 have the same energy density. The higher the slope value, the larger the energy density.Uniform Energy Densityr e w o P r e s a LScanning SpeedFig. 6 Process Window for SLM Raymor Ti-6Al-4V PowderF value99.2611.27The Zone I parameters are capable of building fully dense specimens with Raymor Ti-6Al- 4V powder. It is notable that energy density should be sufficient to avoid incomplete melting. In addition, even if the energy density meets the minimum requirement, porosity could also be included in the specimen. For example, the lower portion of the dashed line in Fig. 6 demonstrates that these laser power and scanning speed cannot be used to fabricate porosity free specimen.

For example, the lower portion of the dashed line in Fig. 6 demonstrates that these laser power and scanning speed cannot be used to fabricate porosity free specimen. Specimens may not be built using Zone OH parameters. Because laser irradiation produces a large quantity of heat which is unable to be conducted away from the specimen immediately. The accumulated strain energy results in serious deformation causing collision with the recoating unit.Zone II and III parameters can be used to fabricate Ti-6Al-4V specimens with porosity. These parameters are known as “marginal parameters” in this study because they are distributed to the margins of Zone I. It is a straightforward situation that insufficient energy density causes smaller melt pools and partially molten powder in Zone III. Thus, voids among powder particles are entrapped under the solidified hatch lines so that the overall density of the specimen is lowered.

Thus, voids among powder particles are entrapped under the solidified hatch lines so that the overall density of the specimen is lowered. However, defects also exist in the specimens fabricated by the Zone II parameters. This implies that extra energy not only causes microscopic structural rearrangements of SLM Ti-6Al-4V, but also a complex transport phenomenon of macroscopic masses over macroscopic distances, with ablation and compression of material and with large internal energies compared to chemical activation energies [13].In order to analyze the melting process under different processing parameters, typical solidified top surface were observed under SEM, as shown in Fig. 7. Each picture corresponds to the processing parameter combination specified in Fig.

Because hatch spacing is constant, the reduced melt pool size results in decreased overlap, which deteriorates the top surface’s roughness, as shown in Fig. 7 (d) and (g). Although voids occasionally appear due to discontinuous hatch lines, they are eliminated by the following recoating and melting process. Thus, no apparent porosity was detectable in the specimens through Archimedes method and destructive characterization techniques. Fig 7 (a), (b), (e) and (i) show the top surface melted by the Zone II marginal parameters. The melt pool was extremely extended with high energy density, especially in the ends of each hatch line due to the ramping phenomenon of scanning speed. On the other hand, energy deficiency causes the melting discontinuity when using the Zone III marginal parameters. As shown in Fig.

On the other hand, energy deficiency causes the melting discontinuity when using the Zone III marginal parameters. As shown in Fig. 7 (h), (j), (k) and (l), many breaking spots are visible along the hatch lines.4.1.3 Defects Caused by Zone II Marginal ParametersA large melt pool results in the increased overlap between hatch lines. However, it can be seen that the highly overlapped hatch lines do not necessarily form voids or defects on the remelted top surface. A certain amount of spherical particles and pits are observable on a solidified top surface, as shown in Fig. 8. It is believed that the porosity of specimens built by the Zone II marginal parameters is directly correlated with these particles and pits. When melting the Ti-6Al- 4V powder at higher energy density, laser irradiation not only causes larger melt pool, but also extra heat which evaporates the molten materials. The melt pool is subject to recoil force from the evolving vapor, which ejects molten materials [14].

The melt pool is subject to recoil force from the evolving vapor, which ejects molten materials [14]. This mechanism is similar to the ablating process of laser drilling. But the materials are even more easily extracted from a powder bed.ParticlePit30 µmFig. 8 Scanning Electron Microscopy of Pits and Particles on Top Surface of SLM SpecimenThe molten metal rapidly solidifies above the building platform and forms spherical particles, some of which fall down and are then welded together with the top surface of the specimens. Small welded particles will be melted together with recoated powder particles. However, large particles, which attach to the top surface, will be easily touched and removed by the recoating blade, as shown in Fig. 9. Pits will be formed with a hollow spherical crown shape after removal. Large pits could be filled with recoated powder. But small pits could be easily entrapped underneath new layers as defects.

9. Pits will be formed with a hollow spherical crown shape after removal. Large pits could be filled with recoated powder. But small pits could be easily entrapped underneath new layers as defects. Due to their occasional formation, these defects aredistributed stochastically in the specimen, instead of directly being correlated with processing parameters. This explains why the uncertainty value of specimen porosity built by the Zone II marginal parameters is larger than that by the Zone III marginal parameters.VVWelded ParticleRecoating bladePitFig. 9 Schematic of Defect Caused by Recoating during SLM Process4.1.4 Defects Caused by Zone III Marginal ParametersThe discontinuity of hatch lines results from the velocity inequality between laser scanning speed and melt pool interface motion. When laser power is applied to the powder bed, the powder particles form a melt pool in a transient process.

When laser power is applied to the powder bed, the powder particles form a melt pool in a transient process. Temperature of the interface between the molten material and un-melted powder is around the melting point (1660 ºC for Ti-6Al-4V). If there is no enough energy furnished to the interface to preserve melting and steady melt pool, the velocity of the interface will be occasionally slower than that of the laser spot motion (i.e. Vmelt pool < Vlaser), as illustrated in Fig. 10 (a). Thus, hatch lines can hardly be formed continuously. Moreover, due to the reduced melt pool size, the overlap between two hatch lines is very limited, which exacerbates the wetting condition for the following layer. If the molten material fails to wet on the previous layer and adjacent hatch line (Fig. 10 (b)), it will easily accumulate to form a rounded surface due to surface tension.

If the molten material fails to wet on the previous layer and adjacent hatch line (Fig. 10 (b)), it will easily accumulate to form a rounded surface due to surface tension. Therefore, the voids or pores are seriously generated.(a) Laser melting metallic powder (b) Top view of two adjacent hatch lines Fig. 10 Schematic of Defect Caused by Zone III Marginal Parameters of SLM ProcessA typical solidified surface is shown in Fig. 11, which was fabricated by the Zone III marginal parameters. The scanning direction is indicated in the figure. Some powder particles were sintered on the top surface which are visible through the voids. The voids are randomly distributed with irregular morphology along and between hatch lines.P = 80W V = 1080mm/sSintered Particles100 µmFig.

The voids are randomly distributed with irregular morphology along and between hatch lines.P = 80W V = 1080mm/sSintered Particles100 µmFig. 11 Top Surface Fabricated by Zone III Marginal Parameters of SLM Process4.2 Effect of Processing Parameters on EBM Ti-6Al-4V 4.2.1 Mean Porosity Distribution and ANOVAEBM specimens’ mean porosity versus individual experimental factor is shown in Fig 12. The porosity for each level of a particular factor corresponds to an average of three experimental runs at that level, typically with four replicates, representing an average of 12 density measurements.6%6%) D R ( y t i s o r o P4%2%) D R ( y t i s o r o P4%2%0%0%01020304005101520Max Current (mA)Focus Offset (mA)6%6%) D R ( y t i s o r o P4%2%) D R ( y t i s o r o P4%2%0%0%0.10.150.20.250.3060120180240Line Offset (mm)Speed Function IndexFig.

12 Mean Porosity at Each Level of Taguchi Experimental FactorsANOVA was also performed to statistically evaluate the effect of each parameter on porosity of EBM Ti-6Al-4V specimens. The results of ANOVA are summarized in Table 4. It can be seen that all F values are larger than F0.05, 2, 27. Thus, all experimental parameters could be considered as significant factors influencing the porosity (95% confidence level). The speedfunction number has the most significant influence on specimens’ porosity, followed by line offset, focus offset, and max current.Table 4 ANOVA of Taguchi ExperimentSourceDegree of freedomSums of Squares (SS) Mean Square (MS) F valueMax Current20.040.0258.94Focus Offset Line Offset Speed Func. No. Error Total F0.05, 2, 27=3.352 2 2 27 350.06 0.06 0.23 0.01 0.400.03 0.03 0.12 0.0093.14 102.02 376.73However, it is difficult to correlate the speed function index with actual energy input based on the existing references.

Thus, line offset, focus offset and max current were studied by using OFAT method with a constant speed function index (default 98 for EBM Ti-6Al-4V powder) at layer thickness 50 µm. The OFAT experiments also illustrate the significance of line offset and focus offset to porosity change. Based on the porosity distribution, the increased line offset and focus offset could also be considered marginal parameters. Otherwise, there is no apparent porosity fabricated by varying max current. It is assumed that the Arcam system is able to regulate the scanning speed automatically according to the powder bed temperature, even at the reduced beam current. Thus, analysis and discussion about EBM processing parameters are mainly focused on line offset and focus offset.4.2.2 Defects Caused by the Increased Line OffsetPorosity distribution versus line offset is shown in Fig. 13.

13. It can be seen that the porosityis increased by increasing line offset, especially when the line offset is larger than 0.18mm.5%4%) D R ( y t i s o r o P3%2%1%0%0.060.10.140.180.220.260.30.341%Line Offset (mm)Fig. 13 Porosity versus Line OffsetLine offset refers to the distance between two hatch lines, similar to hatch spacing in SLM. According to the energy density equation, increased line offset will reduce the energy density value. The OFAT experimental result indicates that porosity is dramatically increased when the line offset value is larger than 0.18mm. In order to correlate line offset with defect generation, solidified hatch lines are shown in Fig. 14.(a) line offset=0.10mm(b) line offset=0.14mm(c) line offset=0.18mm0.5 mm0.5 mm0.5 mm(d) line offset=0.22mm(e) line offset=0.26mm(f) line offset=0.30mm0.5 mm0.5 mm0.5 mmFig.

14 Solidified Top Surfaces of EBM Specimens by Varying Line OffsetWhen line offset is 0.1mm (default value), the solidified hatch lines are arranged equally. No apparent pores or defects are formed between hatch lines. However, overlap between hatch lines will be reduced if the line offset is increased to 0.14 and 0.18mm. Although there are no visible pores or voids on the top surface (Fig. 14 (b) and (c)), a small amount of porosity is detectable from the porosity curve shown in Fig. 7. This could be attributed to the un-melted powder under the adjacent hatch lines, as shown in Fig. 15. The powder gap was trapped inside the specimen as defects. This means not all powder is melted during the EBM process if the line offset value is increased.Line OffsetIncreased Line OffsetLiquid StateSolid StateUn-melted PowderFig. 15 Schematic of Melt Pool with Increased Line Offset during EBM ProcessThe porosity is not obvious until the line offset is increased to 0.22mm.

15 Schematic of Melt Pool with Increased Line Offset during EBM ProcessThe porosity is not obvious until the line offset is increased to 0.22mm. Apparent defects are observable on the surface, as shown in Fig. 14 (d). When line offset is increased to 0.26mm and 0.30mm, the melt pool fails to contact the previous hatch line, which leads to bad wetting conditions. In order to reduce total surface free energy, the melted powder easily gathers together and flows above the un-melted powder, causing voids between hatch lines, as shown in Fig. 14 (d)-(f). The un-melted powder is visible through the pores of the top surface. The dual effects of voids and powder gap seriously increase the porosity of specimens made with a large line offset value. According to the solidified surface, it can be inferred that melt pool width is larger than 0.18mm, but less than 0.22mm.4.2.3 Defects caused by the Increased Focus OffsetPorosity distribution versus focus offset are shown in Fig. 16.

16. The focus offset value doesnot significantly change the porosity until 16 mA.5%4%) D R ( y t i s o r o P3%2%1%0%1%04820 Focus Offset (mA)12162428Fig. 16 Porosity versus Focus OffsetFocus offset is the additional current running through the respective electromagnetic coil and can be translated into an offset of the focal plane from its zero position [15]. Increasing the focus offset value results in a larger beam diameter during the EBM process. This can lead to an increased beam spot, but less energy concentration. Thus, energy density is reduced. Solidified surface was also observed under microscope and shown in Fig. 17.(a) focus offset=4mA(b) focus offset=8mA(c) focus offset=12mAPore0.5 mm0.5 mm0.5 mm(d) focus offset=16mA(e) focus offset=20mA(f) focus offset=24mAPore0.5 mm0.5 mm0.5 mmFig.

17 Solidified Top Surfaces of EBM Specimens by Varying Line OffsetA small focus offset value results in concentrated electron beam spot and profound melting depth. Thus, the penetration depth of an electron beam into the powder layer is greater than that of the laser beam [16]. Previous solidified titanium is also melted to form melt pools which overlap each other to ensure hatching consistency. Each solidified hatch line forms a ridge shape and causes a rough surface. By increasing the focus offset value, the morphology of the melt pool is changed. As shown in Fig. 18, the horizontal melt pool size is enlarged, while the melting depth is reduced.OverlappingOverlappingOverlappingSolid StateUn-melted powderFig. 18 Schematic of Melting Pool with Increased Focus Offset during EBM ProcessIncreased overlap area improves the surface roughness significantly as shown in Fig. 17 (c, d). But small pores may be generated at localized areas.

17 (c, d). But small pores may be generated at localized areas. These spherical pores result from gaseous argon entrapped in the powder particles caused by the atomization process [8]. The melted powder releases the trapped gas which causes the small defects. However, these defects are barely detectable using the Archimedes method. When focus offset value is higher than 16mA, the increased overlap area and reduced melting depth create an unstable melt pool due to underneath un-melted powder, as shown in Fig. 18. The flowability of molten powder causes interstitial shortage which is reflected by the increased porosity.4.3 Comparison of Defect Formation between SLM and EBMAlthough SLM and EBM are similar in principle, defects are generated in different pattern when deviating processing parameters from optimized parameters. In the SLM process, high energy density causes over melting on the powder bed.

In the SLM process, high energy density causes over melting on the powder bed. The defects are resulted from the blade scraping particles, which are solidified from the ejected molten materials on the surface. Thus, the defects generation is mainly attributed to a physical removal process when using Zone II marginal parameters. By contrast, Arcam EBM system adopts a more complicated feedback control design, compared to EOS system. Higher energy density is hard to be achieved during the EBM process. Generally, no defects could be fabricated even though the electron beam power is increased. Beam current and scanning speed are dynamically controlled to maintain a steady melt pool.If the input energy fails to melt the metallic powder completely, the defects will be easily formed due to the voids between the remelted materials. In the SLM process, the Zone III defects are attributed to the breaking spots of hatch lines, or gaps between two hatch lines due to the reduced melt pool size.

In the SLM process, the Zone III defects are attributed to the breaking spots of hatch lines, or gaps between two hatch lines due to the reduced melt pool size. But the increased line offset or focus offset causes the melt pool in an unstable situation in the EBM process, because of the unmelted powder around melt pool. Serious flowability of the molten material results in large voids on the powder bed. Therefore, according to the microscopy, it is anticipated that EBM defects are larger than the SLM defects if energy density is lowered during the melting process.5. ConclusionPorosity distribution was investigated by varying processing parameters in this study. The change of processing parameters is directly related to the energy to the Ti-6Al-4V powder bed. Although energy density has a significant effect on defect generation, it is not linearly correlated with porosity.

Although energy density has a significant effect on defect generation, it is not linearly correlated with porosity. Marginal parameters are responsible for the defect generation during SLM and EBM processes.In the SLM process, splashed particles are generated by the Zone II marginal parameters and then welded to the top surface. The layer deficiency is caused by the recoater removing the large welded particles. Thus, Zone II marginal parameters indirectly generate defects, which are distributed stochastically in SLM specimens. Defects can be directly formed by Zone III marginal parameters. The insufficient energy density results in voids along or between hatch lines.For EBM process, the increased line offset and focus offset are similar to the Zone III marginal parameters of SLM. Defects are formed due to insufficient energy density.

Defects are formed due to insufficient energy density. Interstices between un-melted particles account for a certain amount of porosity in the specimens, while the deteriorated hatch lines exacerbates the inclusion of defects.AcknowledgementsThe authors gratefully acknowledge the support of the Office of Naval Research through awards N00014-09-1-0147 and N00014-10-1-0800, Technical Monitor: Dr. Ignacio Perez. The authors also express their gratitude to the staff of the Rapid Prototyping Center at the University of Louisville for their assistance.Reference [1] W.O. Soboyejo, T.S. Srivatsan. Advanced Structural Materials: Properties, Design Optimization, and Applications. CRC Press 2006, 359-400[2] J. Hiemenz. Electron beam melting. Advanced Materials and Processes. 165(3), 2007, 45-46[3] G. Levy, R. Schindel, J.P. Kruth. Rapid manufacturing and rapid tooling with layer manufacturing (LM) technologies, state of the art and future perspectives.

The laser power, the scan speed, and the number of repeat scans are varied in SLE. These parameters are combined to define a new measure such as the energy density. Microstructures of the samples are investigated using optical microscopy and scanning electron microscopy. Uniaxial tension tests are performed on samples in longitudinal direction and relevant data is extracted. This work is sponsored by the Office of Naval Research through grant N00014-14-1-0658.Keywords: Additive Manufacturing, Scanning Laser Epitaxy (SLE), Nickel-base, Superalloys, CMSX-4®, Mechanical TestingIntroductionIn order to increase the efficiency of gas turbine aero engines, investment cast equiaxed (EQ) turbine component parts are increasingly being replaced with directionally solidified (DS) components that have either columnar or single crystal (SX) microstructures.

However, due to prolonged period of operation, these DS blades suffer from damage and wear limiting their lifetime Such reduction in their effective life necessitates the replacement of the numerous airfoils within the engines. Due to the high cost of producing DS/SX cast turbine blades, the cost of replacing each blade is several thousand dollars. With each engine containing several hundred blades, the cost to replace everyone can reach hundreds of thousands of dollars. Hence, it is of great interest to develop an additive manufacturing (AM)-based process which will restore the SX/DS microstructure at the damaged location allowing the blades to be reused rather than discarded.Scanning laser epitaxy (SLE) is a laser-based manufacturing process for deposition of EQ, DS, and SX structures onto superalloy substrates through the controlled melting and re-solidification of superalloy powders.

In this particular implementation of the process, a laser beam guided by a set of high-speed galvanometer scanners allows for tight control of the amount of energy being applied to the top of the powder bed as well as the speed at which the melt pool moves across the substrate. Under the proper operating conditions and with enough substrate re-melt, the solidification of the powder melt pool will follow the microstructures of the underlying substrate, allowing for directional and even SX growth [1]. The SLE process has shown significant potential for one-step repair of CMSX-4® [2-4], René N5 [5, 6], René 142 [7], MAR-M247 [8], René 80 [9], and IN100 [10, 11].186CMSX-4® was developed through the joint efforts of Canon-Muskegon and Allison in the 1990s [12]. This “second generation” SX superalloy alloy showed an improvement of 35 oC in the turbine airfoil temperature capability compared to CMSX-2® and CMSX-3® [13]. CMSX-4® contains 3 wt.

This “second generation” SX superalloy alloy showed an improvement of 35 oC in the turbine airfoil temperature capability compared to CMSX-2® and CMSX-3® [13]. CMSX-4® contains 3 wt. % of rhenium (Re) similar to other “second generation” SX superalloys. Re was shown to delay coarsening of the phase by partially segregating to the  matrix [14]. CMSX-4® also enjoys solid-solution-strengthening due to the presence of chromium (Cr), tungsten (W), and tantalum (Ta) and precipitation-hardening due to the presence of aluminum (Al) and titanium (Ti) [12].Current work focuses on establishing an optimal operating range for the various process parameters in SLE. The effects of these different processing parameters on the deposit characteristics are examined. The laser power, the scan speed, and the number of repeat scans are varied in SLE. These parameters are combined to define a new measure such as the energy density.

The laser power, the scan speed, and the number of repeat scans are varied in SLE. These parameters are combined to define a new measure such as the energy density. Microstructures of the samples are investigated using optical microscopy and scanning electron microscopy. Uniaxial tension tests are performed on samples in longitudinal direction. The yield strength, ultimate tensile strength, modulus and ductility values are extractedExperimentalThe CMSX-4® powder was produced by Praxair Surface Technologies using an atomization process and had a particle size ranging from 85-150 µm. The SLE process was conducted on rectangular investment-cast SX CMSX-4® substrates having dimensions of 38.10 mm x 11.85 mm x 2.54 mm. Each substrate was placed into a 38.10 mm x 11.85 mm recess cut into an IN625 base plate. The CMSX-4® powder was placed above the substrates using rectangular wells cut into an Aluminum mask plate.Table I. Chemical composition of the CMSX-4® powder (wt.

The CMSX-4® powder was placed above the substrates using rectangular wells cut into an Aluminum mask plate.Table I. Chemical composition of the CMSX-4® powder (wt. %)CMSX-4®Cr Co Mo Re W Al Ti Ta Ni 5.6 1.0 6.5 Bal 6.59.70.4 3.0 6.4After the samples were prepared, they were placed into a Terra Universal atmospheric glove box which was then purged with argon. A 1kW Ytterbium fiber laser (IPG Photonics, Model: YLS- 1000) was used in conjunction with a Cambridge Technologies galvanometer scanner to focus the beam to a diameter of 40 m at the top of the substrate. A consistent raster scan pattern across the width of the sample was used to propagate a linear melt pool along the substrate.In order to facilitate reduced order parameter space, the SLE process parameters namely the laser power, the scan speed, the number of repeat scans, and the scan spacing are combined to define a new measure, the scan energy density (E) [4].

E is defined as follows:(1) tVSSPEPHere, P is the laser power, N is the number of repeat scans, V is the raster scan speed, SS is theraster scan spacing, and tp is the powder thickness and.187Microstructural Investigation and Mechanical TestingEach sample was sectioned along the length and width. Using a Buehler automated disc saw, each sample was first sectioned lengthwise, and then sectioned by multiple widthwise cuts, as shown in Figure 1. Each section was mounted in Bakelite and polished to a mirror finish; starting with 80 grit paper and progressively increasing the size to 1200 grit. The samples were then rough- polished using 5 µm and 3 µm diamond solutions. Finally, the samples were fine polished using a 0.5 µm colloidal alumina suspension.

The samples were then rough- polished using 5 µm and 3 µm diamond solutions. Finally, the samples were fine polished using a 0.5 µm colloidal alumina suspension. The polished samples were then etched with Marble’s reagent (50 ml HCl, 50 ml H2O, and 10.0 gm CuSO4) to eliminate the phase and reveal the dendritic microstructure.A Leica DM6000 optical microscope was thereafter used to take the images. The microstructural investigation of the SLE processed CMSX-4® was carried out on a Hitachi SU8230 SEM. Vickers microhardness measurements were carried out using a Buehler microhardness indenter with a fixed load at 2000 gf. The hardness values were extracted in Vickers microhardness scale (HV). Details on hardness measurements may be found elsewhere [11]. The tensile properties of the CMSX-4® specimens were determined according to ASTM D638 using an Instron 33R 4466 equipped with 10 kN load cell.Figure 1.

The tensile properties of the CMSX-4® specimens were determined according to ASTM D638 using an Instron 33R 4466 equipped with 10 kN load cell.Figure 1. Sections used in metallographic analysisResults and DiscussionAs illustrated in Figure 2 and 3, the longitudinal sections showed a full metallographic bond along the entire length of the samples. The transition from the investment cast CMSX-4® substrate to the SLE-deposited material is indicated by the sharp change in the microstructure such as the dendrite size. The dense deposit is free from porosity, hot tearing or stress cracking. The characteristics of the samples illustrated in Figure 2 and 3 are reported in Table II.Sample 1 was processed with E of 52.73 J/mm3 while Sample 2 was processed with E of 18.75 J/mm3. A previous study on CMSX-4® showed that higher values of E showed lower SX ratio188(defined as the ratio of average total height and the average deposit height) [4].

A previous study on CMSX-4® showed that higher values of E showed lower SX ratio188(defined as the ratio of average total height and the average deposit height) [4]. Following the previous results, the current study also showed that Sample 1 had lower SX ratio.2000 mFigure 2. Representative longitudinal cross-section of the as-deposited Sample 1.This sampleis processed at E = 52.73 J/mm3.2000 mFigure 3. Representative longitudinal cross-section of the as-deposited Sample 2. This sampleis processed at E = 18.75 J/mm3.Table I. Characteristics of Sample 1 and Sample 2Sample #1 2# of Repeats600 400Power (W)750 500Scan Speed (mm/s) 800 750Scan Spacing (m) 12.7 25.4Powder Thickness (mm) 1.4 1.4Energy Density (J/mm3) 52.73 18.75SX Ratio0.39 0.75The orientation of the columnar growth in the SLE-deposited material showed little to no angular misorientation as compared to the underlying substrate as shown in Figure 4(a).

Figure 4(b) illustrates the transition from SX growth with [001] directionality to a shift in orientation. This change is known as oriented to misoriented transition (OMT). On some samples, the top of the deposit suffered from EQ growth. This is known as the columnar to EQ transition (CET).189500 mCET(b)(a)100 mOMTFigure 4. (a) Transition from substrate to deposit and (b) transition regions in the deposit.1.251.301.351.401.451.501.55350375400425450475500SubstrateInterfaceDeposit -[001]Deposit -[100], EQy(Gpa)LocationHardness (HV)Figure 5. Summary of microhardness results for CMSX-4®.In the present study, the [001] columnar region typically showed higher hardness than either the [100] columnar zone or the EQ region. The results indicated a hardness value in the range of 430-500 HV in the deposit region. The substrate region showed hardness values of order 400 HV. Hence, the hardness in the deposit region is about 10% higher than the substrate region.

The substrate region showed hardness values of order 400 HV. Hence, the hardness in the deposit region is about 10% higher than the substrate region. The yield stress was approximated as, y ~ HV/3 (HV measured in GPa) [8].Figure 6 shows the load vs. displacement curves for Sample 1 and Sample 2. Sample 2 showed higher values of modulus, lower values of elongation, and higher yield strength as compared to Sample 1 as illustrated in Figure 6. This difference is due to the fact that Sample 2 had higher SX190ratio compared to Sample 1 and therefore, it has higher microhardness. It is noted that the results are presented as load-dispalcement instead of stress-strain as the cross-section (thickness of the samples was different for the two samples and was varying along the length within the same sample).

That is the reason that no absolute values of the modulus or strength are mentioned.05001000150020002500300001234Sample 2Displacement (mm)Load (N)Sample 1Figure 6: Load vs displacement curves for samples 1 and 2 based on tensile testingConclusionsIn the present study, two different samples deposited by SLE at two different energy densities were explored. The sample produced at higher values of E showed lower SX ratio and therefore, lower yield strength. It is concluded from the present research that higher SX ratio is desirable for the purpose of obtaining higher yield strength and modulus.Acknowledgments and DisclosuresThis work is sponsored by the Office of Naval Research through grant N00014-14-1-0658. Dr. Suman Das is a cofounder of DDM Systems, a start-up company commercializing SLE technology. Dr. Das and Georgia Tech are entitled to royalties derived from DDM Systems’ sale of products related to the research described in this paper.

Dr. Das and Georgia Tech are entitled to royalties derived from DDM Systems’ sale of products related to the research described in this paper. This study could affect their personal financial status. The terms of this arrangement have been reviewed and approved by Georgia Tech in accordance with its conflict of interest policies.References[1] Basak A, Das S. Epitaxy and Microstructure Evolution in Metal Additive Manufacturing. Annual Review of Materials Research 2016;46:125-49. [2] Acharya R, Bansal R, Gambone JJ, Das S. A Coupled Thermal, Fluid Flow, and Solidification Model for the Processing of Single-Crystal Alloy CMSX-4 Through Scanning Laser Epitaxy for191Turbine Engine Hot-Section Component Repair (Part I). Metallurgical and Materials Transactions B 2014;45:2247-61. [3] Acharya R, Bansal R, Gambone JJ, Das S.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEFFECTS OF SCANNING STRATEGY ON RESIDUAL STRESS FORMATION IN ADDITIVELY MANUFACTURED TI-6AL-4V PARTSMohammad Masoomi1, Scott M. Thompson1,†, Nima Shamsaei1, Meysam Haghshenas21 Laboratory for Fatigue & Additive Manufacturing Excellence, Department of Mechanical Engineering, Auburn University, Auburn, AL, USA 368492 Mechanical Engineering Department, University of North Dakota, Grand Forks, ND, USA 58202† Corresponding author: Email: smthompson@auburn.edu Phone: (334) 844 4867AbstractParts fabricated via directed energy additive manufacturing (AM) can experience very high, localized temperature gradients during manufacture. These temperature gradients are conducive to the formation of a complex residual stress field within such parts.

These temperature gradients are conducive to the formation of a complex residual stress field within such parts. In the study, a thermo-mechanical model is employed for predicting the temperature distribution and residual stress in Ti-6Al-4V parts fabricated using laser- powder bed fusion (L-PBF). The result is utilized for determining a relationship between local part temperature gradients with generated residual stress. Using this numerical model, the effects of scan patterns are investigated.IntroductionLaser-Powder Bed Fusion (L-PBF) is an additive manufacturing (AM) method used for producing dense, metallic parts by way of melting successive layers of powder material by a laser. One of the challenges facing this method is residual stress formation and deformation of built parts [1,2]. Temperature gradients and cooling rates experienced during L-PBF can result in adverse material side effects such as residual stress or pore/defect formation, respectively [3,4].

Temperature gradients and cooling rates experienced during L-PBF can result in adverse material side effects such as residual stress or pore/defect formation, respectively [3,4]. There are different methods to evaluate the residual stress of the parts. In this study, neutron diffraction was employed to measure internal residual stresses at various locations along stainless steel (SS) 17-4 PH specimens additively manufactured via laser-powder bed fusion (L-PBF). Of these specimens, two were rods (diameter = 8 mm, length = 80 mm) built vertically upward and one a parallelepiped (8 × 80 × 9 mm3) built with its longest edge parallel to ground. Simulations were performed using the commercially- available software COMSOL Multiphysics® 5.1 in conjunction with custom coding scripts.Approach and Numerical ModelDuring the fabrication of part, powder will melt and solidify. To model transient heat transfer,numerical model was developed.

To model transient heat transfer,numerical model was developed. The heat transfer in the part is governed by the heat equation:𝜌𝑐i(Ti)|𝑡o−∆𝑡𝜕Ti(𝑥, 𝑦, 𝑧, 𝑡) ∂𝑡| 𝑡o= ∇𝑘i(Ti)|𝑡o−∆𝑡∇Ti(𝑥, 𝑦, 𝑧, 𝑡)|𝑡o(1)2546where ci is specific heat capacity and ki is continuum thermal conductivity and ρ is the density. All modes of heat transfer, i.e. conduction, convection and thermal radiation, are taken into account in the model. The numerical model used in the study is developed by authors and explained in details at [5,6].Since the local temperature gradient consists of three spatial components, i.e. parallel-to-track (x), perpendicular-to-track (y) and normal-with-substrate (z) directions, the magnitude of local temperature gradients was calculated herein.

Process parameters used for the experimental studySubstrate material Substrate size Powder description Powder layer thickness Powder bed porosity Laser spot diameter Laser power Laser wavelength Scan speed Powder bed absorptance Absorption coefficient Emissivity Melt pool liquid thermal conductivity Shielding gas type Shielding gas temperature Shielding gas flow rate Chamber wall temperature Substrate temperature Convection heat transfer coefficientTi-6Al-4V 10 x 10 x 1 cm3 Gas-atomized, air-dried 30 μm 0.4 70 μm 48 W 1075 nm 300 mm/s 0.25 [7] 106 cm-1 [8] 0.35 [9] 25 – 45 W/m∙K Argon 20 ºC 167 cm3/s 20 ºC 20 ºC 12.7 W/m2∙KNumerical results and discussionThe size of the sample numerically investigated was 10 × 1 × 0.03 mm3 and the PBF of the layer atop a previously-melted layer was simulated The tracks were unidirectional, meaning they went from right to left. The temperature and temperature gradient at the beginning, middle and end of a track are presented in Figure 1.

The temperature and temperature gradient at the beginning, middle and end of a track are presented in Figure 1. It can be observed that temperature gradient decreases along the track.2547〉‖(2)Figure 1. Temperature and temperature gradient at (a) start of track, (b) middle of track and (c) end of track.The reason for lower temperature at the end of track can be attributed to heat accumulation effects. Since temperature gradients drive the formation of thermal stresses, it is reasonable to assume that higher temperature gradients will result in higher residual stresses and vice versa. Parry et al. [10] showed that residual stress decreases toward the end of a track during PBF. It is also important to investigate the effects of alternate scan strategy, instead of the unidirectional scan strategy. It is already presented that temperature gradient is highest at the start of track and will decrease significantly along the track.

It is already presented that temperature gradient is highest at the start of track and will decrease significantly along the track. Using an alternate scan strategy, both ends would experience higher temperature gradient compared to middle of track. Based on these results, it is expected that temperature gradients would be higher near the surfaces of parts. Masoomi et al. [11] presented experimental residual stress data gathered from neutron diffraction.2548Based on these data, when an alternating scan strategy is used, residual stress will be much higher near the surface as compared to the residual stress near the center.ConclusionsIn the study, the effects of thermal response on residual stress were studied. A framework to correlate thermal response to residual stress was provided. Effects of scan strategy on residual stress were studied.

A framework to correlate thermal response to residual stress was provided. Effects of scan strategy on residual stress were studied. It was shown that temperature gradients and consequently residual stress will be lower at the end of a laser track relative to the beginning of it. It was concluded that using an alternating scan strategy will cause residual stress to be lowest at the center of parts and highest at both ends.AcknowledgementsThis work has been supported in part by NSF Grant #1657195.References [1] K. Dai, L. Shaw, Distortion minimization of laser-processed components through control of laserscanning patterns, Rapid Prototyp. J. 8 (2002) 270–276.[2] B. Vrancken, V. Cain, R. Knutsen, J. Van Humbeeck, Residual stress via the contour method incompact tension specimens produced via selective laser melting, Scr. Mater. 87 (2014) 29–32. doi:10.1016/j.scriptamat.2014.05.016.[3] C. Weingarten, D. Buchbinder, N. Pirch, W. Meiners, K. Wissenbach, R.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEFFECTS OF SPATIAL ENERGY DISTRIBUTION ON DEFECTS AND FRACTURE OF LPBF 316L STAINLESS STEELElliott Jost*,†, John Miers*,†, Aron Robinson†, David Moore†, Christopher Saldana*George W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332 †Sandia National Laboratories, Albuquerque, NM 87123AbstractMeasures of energy input and spatial energy distribution during laser powder bed fusionadditive manufacturing have significant implications for the build quality of parts, specifically relating to formation of internal defects during processing.

In this study, scanning electron microscopy was leveraged to investigate the effects of these distributions on the mechanical performance of parts manufactured using laser powder bed fusion as seen through the fracture surfaces resulting from uniaxial tensile testing. Variation in spatial energy density is shown to manifest in differences in defect morphology and mechanical properties. Computed tomography and scanning electron microscopy inspections revealed significant evidence of porosity acting as failure mechanisms in printed parts. These results establish an improved understanding of the effects of spatial energy distributions in laser powder bed fusion on mechanical performance.IntroductionAdditive manufacturing (AM), the layer-by-layer buildup of a 3D component based on a digital model, offers significant advantages over traditional manufacturing techniques in terms of cost-to-complexity ratio as well as time and cost for small batches of parts [1].

However, along with these advantages comes the difficulty of safely implementing AM components in high- consequence applications, such as for rotating engine blades where the internal defect structure of AM components must be tightly controlled. Consequently, qualification, the determination of the integrity of a component, has become an increasingly important subject of interest in the research community [2–5]. In-situ monitoring and computed tomography are used to determine part integrity in-situ (during the build process) and post-build, respectively [5].In the present study, comparisons were made between measures of global energy density(GED), mechanical properties, and quantitative scanning electron microscopy fractography measurements of 316L stainless steel tensile samples printed using three separate GED conditions.

The goal of this study is to elucidate the effects of processing defects as they relate to the mechanical properties of laser powder bed fusion (LPBF) parts printed using various laser powers at constant laser speed and hatch spacing.QualificationIn-situ monitoring of the printing process is often used in qualification efforts in large part due to its potential to determine defect formation during a build and its potential use for implementing feedback control of the AM process. In-situ corrections of process parameters can440be used to create so-called “born qualified” parts that are determined to meet a final specification as soon as they are removed from the build plate [6]. Craeghs et al. [7] successfully demonstrated the use of optical sensors to implement feedback control during the printing process for surface roughness improvement.

Craeghs et al. [7] successfully demonstrated the use of optical sensors to implement feedback control during the printing process for surface roughness improvement. In their review, Spears and Gold [8] identified over 50 process parameters relating to the laser, stock powder, powder bed properties, recoating, and build environment that influence the final part quality for a variety of metal material systems, highlighting the difficulty of using in-situ monitoring for qualification.In addition to in-situ monitoring, microcomputed tomography (CT) has been an effectiveresearch tool for the inspection and characterization of AM parts due to its ability to nondestructively take high resolution 2D images and reconstruct them to create 3D volumetric data. This data can be used to visualize defect location and morphology within AM-produced components.

This data can be used to visualize defect location and morphology within AM-produced components. Other methods of inspection such as serial sectioning [9] can be used to inspect parts, but at the expense of a significant time investment and the destruction of the sample. Kim et al. [10] developed 3D image processing techniques to improve the reliability of CT data analysis and also attempted to standardize these methods using CoCr parts. In another study, Kim et al. [11] used CT to directly inspect internal defects of Inconel 625 AM parts. In an extensive review on the use of CT for AM inspection, Thompson et al. [12] discusses the advantages of using CT inspection to measure pore morphology, location, and distribution. Synchrotron CT has also been employed to quantify processing defects in AM parts on the sub- micron scale [13].

Synchrotron CT has also been employed to quantify processing defects in AM parts on the sub- micron scale [13]. Moreover, several studies have utilized CT to understand the conditions under which defects are created across a variety of material systems including 316L, Ti-6Al-4V, and 304L stainless steel [14–16].Correlation of Defects and Mechanical PropertiesTo qualify a component, one must first understand the effects of defects and features ofinterest and the degree to which these defects and features will impact part performance, which is the motivation of this study. Only then can a determination be made as to the integrity of a part. A variety of studies have been performed in order to elucidate the complex relationship between defects and performance in AM parts across a variety of metals [17–22]. However, few of these studies consider or provide quantitative metrics on porosity morphology in attempt to relate defects to properties.

However, few of these studies consider or provide quantitative metrics on porosity morphology in attempt to relate defects to properties. Recently, efforts have been made to understand how pore morphology data can be used to understand the role of defects as stress concentrators [23].In a study on the effects of non-optimum energy density in Ti-6Al-4V AM components,Gong et al. [17] concluded that low-energy density input was more detrimental to part performance than high-energy density input. Madison et al. [18] found that there was little correlation between metrics such as total defect volume or number of defects and properties in 304L stainless steel. Carlton et al. [19], using synchrotron radiography, found that voids can act as crack initiation sites and that effects were most significant in samples with nonuniform defect distributions in 316L.

Carlton et al. [19], using synchrotron radiography, found that voids can act as crack initiation sites and that effects were most significant in samples with nonuniform defect distributions in 316L. It is important to note that while the above studies cover a variety of materials and, as such, their findings cannot be directly applied to 316L stainless steel, their collective conclusions are helpful for understanding the role of defects AM components in general.

Through correlation of pore morphology metrics obtained using SEM images and part441mechanical properties, this paper aims to provide insight into the defect-property relationship in efforts to standardize the qualification of AM components.MethodsManufactureIn the present study, three tensile test sample arrays (75 total samples), CT scans ofwhich are presented in Figure 1a, were printed with an EOS M 290 LPBF machine equipped with a 400W IPG Photonics Laser using 316L stainless steel powder with a nominal particle size distribution of 30-60 μm.15mm50mm6mmFigure 1. a) CT images of As-Built Sample Array b) Photo of the Nominal Build Plate with Analyzed Set Highlighted. Wiper direction is top-to-bottom and Argon flow is from right-to-left.Three build plates containing a variety of parts were printed at three different globalenergy distribution (GED) levels that will be referred to as high, nominal, and low GED levels.

One sample array set was analyzed from each build plate, the build plate location of which is highlighted in Figure 1b. GED, a measure of laser energy input per area [1], can be calculated using Eq. 1, below,(cid:1833)(cid:1831)(cid:1830) =(cid:1842) (cid:1834) (cid:1845)where P is laser power, H is hatch spacing, and S is laser speed. GED is the spatial equivalent of volumetric energy density [1]. In this study, GED was altered by changing the laser power while maintaining constant hatch spacing and laser speed values. GED parameters used in this study are shown in Table 1.442(1)High GEDTable 1. Global Energy Distribution Parameters S (mm/s) H (mm)P (W)2200.091083GED (J/mm2) 2.26Nominal GED1950.0910832.00Low GED1700.0910831.74Computed Tomography InspectionPrior to mechanical testing, all samples were nondestructively inspected using a Nikon M2 Dual Head 225/450 kV computed tomography system.

Samples were inspected at 440 kV and 227 μA while using a 1 mm Cu filter to mitigate beam hardening image artifacts caused by the naturally polychromatic energy output of laboratory x-ray sources. Helical scanning was employed to effectively remove some types of image artifacts while allowing single-scan data acquisition of each of the elongated samples used in this study. To mitigate the interference of circular artifacts with data processing techniques, an off-axis center of rotation was used, which can create higher resolution scans but is prone to introducing additional noise and streaking. A resolution of ~10 μm/voxel was achieved using a Perkin Elmer detector and 2294 projections with 8 frames averaged per projection. Total CT inspection time was approximately 2 hours per array.

A resolution of ~10 μm/voxel was achieved using a Perkin Elmer detector and 2294 projections with 8 frames averaged per projection. Total CT inspection time was approximately 2 hours per array. Reconstruction and porosity analysis were performed using Nikon Metrology X-Tek CT Pro 3D and Volume Graphics VGSTUDIO MAX 3.2™, respectively.The scanned sample arrays were initially segmented using an ISO-50 threshold followedby an advanced surface refinement using a deformable surface technique for a more accurate determination of surface contours. Additionally, noise particles and voids were removed in this process. Following segmentation, planes were fit to the top, right, and front surfaces of each array.

Additionally, noise particles and voids were removed in this process. Following segmentation, planes were fit to the top, right, and front surfaces of each array. Each volume was then digitally registered such that the top plane normal vector is used as the direction for the positive z-axis, the line formed by the intersection of the top and right planes is used for the direction of the x-axis, and the point where all three fit planes intersect is set as the origin of the coordinate system. Plane fit points and the defined coordinate system are shown in green and red in Figure 2a, respectively.Figure 2. Example fit points (colored dots) used to digitally register each object volume. a.) Fit points shown in green on volume. b.) Opaque gage regions with 85% transparent sample array.443The gage region of each tensile sample in each array was isolated for individual porosity assessment.

b.) Opaque gage regions with 85% transparent sample array.443The gage region of each tensile sample in each array was isolated for individual porosity assessment. Since the as-designed dimensions and distribution of the specimens was the same across all arrays examined in the present work, the coordinate system defined above allows for the use of uniform parameters in the definition of the regions of interest for consistent comparison between datasets. To separate the gage region of each tensile sample, a rectangular cuboid region of interest (ROI) was created. A Boolean intersection of this ROI with the segmented material was then performed to obtain a single unified ROI that encompasses the exact contours of the gage regions. Gage regions of the sample array are visualized as opaque in Figure 2b.Finally, porosity in each gage region was identified using VGDefX™, a built-in porosity analysis tool of VGSTUDIO MAX 3.2.

Gage regions of the sample array are visualized as opaque in Figure 2b.Finally, porosity in each gage region was identified using VGDefX™, a built-in porosity analysis tool of VGSTUDIO MAX 3.2. A medium adaptive noise reduction was used along with a 2-voxel offset from the determined surface to clean up the material inside the gage region. Results of the porosity analysis were filtered to only include pores with realistic void shape. Table 2 contains the filter parameters that the identified porosity had to fall within to be considered real. The probability threshold does not represent a percentage, but rather a numerical value assigned to each pore that is the assessment of an internal weighted function in VGDefX™. Compactness is the ratio of pore’s volume to that of a circumscribed sphere of the same diameter. Similarly, sphericity is the ratio of a pore’s surface to that of its circumscribed sphere.Table 2.

Compactness is the ratio of pore’s volume to that of a circumscribed sphere of the same diameter. Similarly, sphericity is the ratio of a pore’s surface to that of its circumscribed sphere.Table 2. Parameters used to filter the results of the porosity analysis in VGDefX™ of VGSTUDIO MAX 3.2.Probability Threshold >0.85Effective Pore Diameter [mm] 0.0625-1.00Compactness0.08-1.00Sphericity0.13-0.65Through CT inspection, it was found that the high GED samples contained high individual pore volume with a moderate number of pores within the gage volume, nominal GED samples had small individual pore volume and a low number of pores, and low GED samples contained small individual pore volume with a large number of pores. Most importantly, porosity distributions were found to be spatially uniform along the z-axis of the gage region, suggesting that fracture surface SEM inspections fairly sample the porosity distributions.

CT porosity measurements were found to agree with SEM fracture surface inspection measurements. Quantitative porosity metrics will be presented later in this paper.High-Throughput Mechanical TestingAn automated high-throughput tensile testing procedure was developed to test tensile bars for rapid characterization of mechanical properties of AM samples. Tensile testing was accomplished using an MTS servo-hydraulic load frame equipped with an Interface 2-kip load cell and a mounting stage for sample movement. Sample were moved using screw-driven linear motion with a KeLing Technologies, Inc. stepper motor and a Temposonics® position sensor. Real-time DIC strain tracking using VIC-Gauge™ software by Correlated Solutions allowed for further automation by using non-contact strain measurements to reduce the need for operator intervention. Two cameras, normal to the front and right faces of each sample, were used for strain tracking cross sectional area measurements.

Two cameras, normal to the front and right faces of each sample, were used for strain tracking cross sectional area measurements. A pixel-to-millimeter ratio was established444using a reference image of known dimensions and applied to the first photo taken of every sample to determine the initial area using a MATLAB algorithm. Samples were speckled with a white undercoat and black speckle pattern via spray paint and pulled at a constant rate of 0.03 mm/s. A MATLAB code was written to process the data and create stress-strain curves as well as generate a results file listing the yield strength at 0.2% offset, ultimate tensile strength, uniform elongation, and the strain to failure for each individual specimen.Scanning Electron Microscopy InspectionFollowing tensile testing, a Zeiss Ultra 60 Field Emission Scanning Electron Microscope (SEM) equipped with a Schottky FEG gun was used to image the 150 fracture surfaces resulting from tensile testing.

Samples were imaged at a working distance of approximately 8 mm and an acceleration voltage of 15 kV using the secondary electron detector to accurately capture topographical information of the fracture surfaces. Samples 1, 5, 13, 21, and 25 were selected as samples of interest from each sample array, as seen in Figure 3, and subjected to closer SEM inspection to determine if fractographic differences existed between samples produced on the edges and middle of the samples, as detailed below.15132125Figure 3. Post-fracture image of sample array. Samples 1 to 25 shown left to right. Dimensions shown in Figure 1a.Quantitative FractographyFIJI, an ImageJ analysis package [24], was used to quantitatively inspect porosity observed on the fracture surfaces of the analyzed samples. The samples of interest were analyzed quantitatively from each sample array using an SEM image containing the entire fracture surface at ~250X magnification and a resolution of ~3μm/pixel.

The samples of interest were analyzed quantitatively from each sample array using an SEM image containing the entire fracture surface at ~250X magnification and a resolution of ~3μm/pixel. Using the freehand selection tool, 647 pores were manually identified across these 15 samples by tracing the outside edge of each pore. Only bottom fracture surfaces, or those still attached to the array as seen in Figure 3 were analyzed quantitatively, as the complimentary fracture surfaces were assumed to be identical for the purposes of this study. Porosity metrics calculated from image analysis include pore area, Feret (longest) diameter, centroid, circularity, as well as distance between pore centroid and fracture surface centroid. An example of pore identification can be seen in Figure 4.445abFigure 4. a) Original SEM image of N13. b) SEM with analyzed pores highlighted in green.A variety of sizes and shapes of pores were identified on the fracture surfaces.

a) Original SEM image of N13. b) SEM with analyzed pores highlighted in green.A variety of sizes and shapes of pores were identified on the fracture surfaces. As willlater be shown quantitatively, the largest pores tended to cluster around the edges of the samples, as exemplified in Figure 4 and Figure 5. Figure 5 shows several of the analyzed samples in a visual summary of the porosity found during the quantitative fractography analysis where analyzed pores are segmented in green. In Figure 5 it can be seen that defects observed by SEM are larger in the high GED samples and significantly smaller, on average, in the low GED samples. Because the focus of this study is internal porosity and not the tortuous surface “crust” heterogeneities detailed in other studies [25], only pores that were determined to be completely internal as seen from SEM imaging were included for analysis.Sample 1Sample 13Sample 25D E G h g i HD E Gl a n i m o ND E G w o LFigure 5.

Visual Summary of Select Analyzed Samples.446Results and DiscussionComparison of Porosity Metrics to Energy DensityFrom the collected quantitative fractography data, comparisons were made between porosity metrics and GED to understand the effect that energy input has on porosity formation in 316L stainless steel AM components. Figure 6 shows comparisons of mean pore area and GED, which reveals a clear trend where high, nominal, and low GED samples have mean areas of 140.2 μm, 67.5 μm, and 32.6 μm, respectively. Figure 6b shows the mean area of the five largest pores on each inspected sample and shows a similar trend, with high, nominal, and low pore sizes of 800.9 μm, 400.4 μm, and 144.6 μm. The five largest pores seen on each fracture surface are considered in this and other figures to understand whether a minimum critical defect size that impacts mechanical performance can be determined for establishing inspection specifications.

Defect size distributions were found to be non-Gaussian due to the extreme outliers in high and nominal GED samples and thus trends should not scale linearly as a result of Gaussian distribution sampling properties but rather due to the dominating effects of the largest defects.150.0~E ::L ., "' 50.0 ~100.00.0a140.2IHigh67.5■NominalGED1000.0~ 200.0 Low~ 200.0 0.0~ 200.0 b~E ::L 600.0 ., "' 400.032.6800.9400.4■NominalGED Low144.6Figure 6. a) Average Sample Pore Area vs. GED. b) 5 Largest Pores Mean Area vs. GED.In addition to pore area, distance between each pore’s 2D centroid and the 2D centroid of the fracture surface, defined as the center of the gage cross-section, was calculated. These results are presented in Figure 7.

These results are presented in Figure 7. As shown, the average distance of the pore from the part center is greatest in the high GED samples and smallest in the low GED samples, indicating that pores exist farther from the center in high GED samples, regardless of size. This trend is even more severe when only considering the five largest pores from each sample where the average distance of the largest five pores is more than four times larger in the high GED than low GED samples. Due to necking behavior observed in the samples, the maximum possible distance is less than 375 μm, half the nominal specimen width, and is estimated to be approximately 325 μm.a233.6.,219.0200.0 I I250.0u C: ~ 150.0 i5 °E 100.0 ., ::L ~ - 50.0 ... ., 0.0 ~HighNominalGED188.6ILow150.0E ., E, 100.0 u C fl ., 0 "" ro v :> b <(50.0V,0.0140.2IHigh67.5■NominalGED Low32.0Figure 7. a) Pore Center to Part Center Distance vs. GED.

a) Pore Center to Part Center Distance vs. GED. b) Biggest 5 Pores Pore Center to Part Center Distance vs. GED447Figure 8 provides a visual representation of pore location and pore distance to part center and shows the 2D location of each pore for each GED type, where location (0,0) is the center of each part and the black rectangle represents the nominal edges of the sample. The reader should note that the circles in Figure 8 are shown for the purposes of representing the location and relative size of the pores using the pore area calculated during SEM inspection and should not be interpreted as an absolute representation of pore size or shape. Figure 8d shows the locations of pores with areas in at least the 95th percentile, most of which occur near the edge of the part. Note that in the distribution of the 33 pores in this percentile range, 21 are in the high GED samples, 11 in the nominal GED, and 1 in the low GED.

Note that in the distribution of the 33 pores in this percentile range, 21 are in the high GED samples, 11 in the nominal GED, and 1 in the low GED. The authors note that pores shown in Figure 8d are primarily concentrated in the upper part of the plot and hypothesize that the cause of this may be due to the part location on the build plate as well as the flow of the inert gas in the build chamber. Further investigation is needed to draw more concrete conclusions.High GED Pores 500 - - -~ - - - - - -~Nominal GED Pores500 - - - - - - - - - - -E :::l..0E :::l..0a500 ~ - - - - -~ - - - - -~ 5005000 µmLow GED Pores500 ~ - - - - - - - - - - - -b500 ~ - - - - -~ - - - - -~ 500 0 µ m > 95th Percentile Pores 500 ~ - - - - - - - - - -~500000 ~~E :::l..0E :::l..0000 000o High o Norn o Low0500 ~ - - - -~ - - - - -~ 500500 ~ - - - -~ - - - - -~ 5005005000 µ m0 µ mc Figure 8. Pore Location Maps.

Pore Location Maps. a) High b) Nominal c) Low GED Condition Pores. d) Pores in 95th percentile or above volumetrically.d448Based on the results displayed in Figure 6, Figure 7, and Figure 8, it is hypothesized thatthe cause of the increase in the number of large pores as well as their location near the edge of the specimens is due to the path of the laser. It is well-established that the likelihood of keyhole porosity formation is increased in laser welding and SLS of stainless steels when energy input is elevated [14–16]. It is plausible that the raster pattern used in the printing of these samples resulted in a non-optimal energy input at the points at which the laser reverses directions on the sample. The instantaneous zero-velocity of the laser can result in a momentary elevation of energy input and cause the formation of keyhole porosity.

The instantaneous zero-velocity of the laser can result in a momentary elevation of energy input and cause the formation of keyhole porosity. In the nominal GED samples, perturbations in laser power and control parameters could have resulted in an increased likelihood of either keyhole porosity formation or lack of fusion around the edges of the samples. In the high GED samples, the increased laser power in addition to the parameter perturbations likely caused the frequent keyhole porosity observed in the SEM inspections of the fracture surfaces.Conversely, the low GED samples did not have sufficient energy input to cause keyholing to occur. Keyhole formation has been linked to normalized enthalpy thresholds, above which keyhole mode melt pool formation dominates over the preferred conduction mode [14]. It is hypothesized that the low GED samples did not exceed the normalized enthalpy threshold for 316L stainless steel, and thus did not experience keyhole porosity formation.

It is hypothesized that the low GED samples did not exceed the normalized enthalpy threshold for 316L stainless steel, and thus did not experience keyhole porosity formation. Gas porosity, however, formed by the entrapment of inert gas flowing in the build chamber, is commonly small and spherical in shape, as seen in the low GED samples. Shape of porosity can be used to distinguish between gas pores and lack of fusion pores, which are common when there is insufficient energy input to completely fuse powders together, resulting in a random pore shape rather than a spherical pore shape.

Due to the low number of large pores in Low GED samples, it is hypothesized that the small pores observed in these samples as well as their relatively uniform spatial distribution across the examined fracture surfaces of all samples are likely gas and lack of fusion porosity.Comparison of Mechanical Properties to Porosity MetricsComparisons were made between the mechanical properties of the and porosity metrics of the SEM-inspected samples to elucidate the complex relationship between defects and final mechanical properties. Several mechanical properties were measured from tensile test data including ultimate tensile strength (UTS), ductility, and yield stress.

Several mechanical properties were measured from tensile test data including ultimate tensile strength (UTS), ductility, and yield stress. Average measured mechanical properties of all 75 samples are shown compared to GED in Figure 9 and to wrought and annealed 316L stainless steel properties in Table 3.495.0 IQ 490.0 ~ 485.0 ;;,- 480.0 ~ 475.0 470.0490.3I I481.0~<ia"~~,f'~~493.7Io.i- v80.0 i 60.0 ? 40.0 .., u ,5 20.00.073.670.9370.060.6 I I 360.0 IIQ a. 350.0 ~ 340.0 £ 330.0 ~ 320.0 .b <I) lJ ~~~<ia":cs~ ~,f'0~ v>336.9I363.6 366.1I I~<ia"~~~~oo.i- vaGEDbGEDCGEDFigure 9. Mechanical Properties of all Samples vs. GED. a) UTS. b) Ductility. c) Yield Strength.449Table 3.

Mechanical Properties of all Samples vs. GED. a) UTS. b) Ductility. c) Yield Strength.449Table 3. Summary of Analyzed Sample Properties of all Samples ± 1 Standard Deviation and Wrought and Annealed 316L Properties [26] UTS (MPa) 490.3 ± 12.34 481.0 ± 9.66 493.7 ± 7.28 515Ductility (%) 60.6 ± 3.99 70.9 ± 4.62 73.6 ± 4.53 60Yield Strength (MPa) 336.9 ± 20.96 363.6 ± 7.66 366.1 ± 12.90 205High GED Nominal GED Low GED Wrought + AnnealedAs can be seen in Figure 9, mechanical properties can vary significantly with GED type. Most significantly, ductility was found to vary greatly with increased GED, as evidenced by the poor performance of the high GED samples at 60.6% compared to 73.6% in the low GED samples, as shown in Figure 9b. UTS and yield strength also varied with GED type, with high samples performing well for UTS at 490.3 MPa and poorly for yield strength at 336.9 MPa.

UTS and yield strength also varied with GED type, with high samples performing well for UTS at 490.3 MPa and poorly for yield strength at 336.9 MPa. However, most important to consider is how these properties vary based on defect metrics such that a minimum critical defect size can be established. As seen in Figure 10a and Figure 10b, pore area seems to have no significant correlation with ultimate tensile strength based on the wide spread of intra-array UTS performance, even in the nominal GED samples.520.00"' ~ 500.00~ 480.00 ::> 460.00.. '••• • • • ••520.00 "' ~ 500.00 ~ 480.00 ::>460.00I• •High GEDNorn GED0.050.0100.0 150.0200.0 250.00.0500.0 1000.0 1500.0Low GEDaAverage Pore Area (µm 2)bAverage Pore Areas (µm 2)Figure 10. UTS vs. Pore Area. a) All Pores.

UTS vs. Pore Area. a) All Pores. b) Top 5 Largest Pores Per SampleFigure 11a and Figure 11b reveal that pore area plays a significant role in the ductility ofsamples by showing a negative linear trend between ductility and average pore area. Perhaps most interestingly, ductility seems to be uniformly affected by mean pore area as well as the area of the 5 largest pores. The persistence of this trend when only considering the largest pores suggests that the largest pores dominate effects on performance and that it may be sufficient to only consider these pores for component inspection and qualification. Because the largest pores on the fracture surface seem to dominate ductility performance, pore size thresholds can be intelligently determined to inform the inspection process.Furthermore, there exists a negative trend between ductility and pore distance to center as shown in Figure 11c.

Because necking occurred less in the low ductility samples, it is difficult to determine whether the observed trend is due to the natural movement inward of the pores during the necking process or whether there is a true correlation between ductility and pore location. Pre-mechanical testing inspection such as computed tomography could be used to avoid this necking effect on inspection.450_ 100.00 ff 50.00~I ••~ 100.00 l f ..,, .. .I 50.00 ·5. .•••:::,0.00 ~-~--~-~-~~-~ 50.0 100.0 150.0 200.0 250.00.0C)0.000.0I500.01000.01500.0Average Pore Area (µm2)Average Pore Area (µm2)ab100.00~ 80.00 >- 60.00 :t:: ..., ·- u ::, a40.0020.00•••___ I_ I _ __, • • • • • •••e High GEDe NomGED0.00e Low GED150.00170.00190.00210.00230.00250.00270.00Average Pore Distance to Center (µm)CFigure 11. Ductility vs. Pore Morphology. a) Pore Area. b) Top 5 Pore Area.

ti; ~ 3:! - Q) >=300.000.0 500.01000.Cl500.0High GEDNorn GEDLow GEDaAverage Pore Area (µm2)bAverage Pore Area (µm2)Figure 12. Yield Strength vs. Pore Area. a) All Pores. b) Top 5 Pores.The dominating effect of the largest pores of a fracture surface on the mechanical performance of a component in terms of ductility and yield strength are significant findings for the progress towards AM part qualification. Relationships presented in this study suggest that defect impact on mechanical performance is governed only by defects in excess of some critical size.

Relationships presented in this study suggest that defect impact on mechanical performance is governed only by defects in excess of some critical size. The determination of minimum critical defect sizes for various loading applications will result in establishment of inspection routines that are capable of adequately inspecting components for critical flaws as well as optimized for time, cost, and complexity of inspection.Although the data presented in this study indicates that only the largest pores at thefracture surface matter, it is unclear, due to the geometrically-identical specimens tested, how451these effects may vary with part size, geometry, or loading condition. Thus, it is unclear whether inspection criteria should be established based on a universal minimum defect size or whether the pore size-to-stressed area ratio is the determining factor in influencing part performance.

Further studies will need to be performed to make the influence of these competing factors clear for the eventual establishment of inspection criteria for AM components.ConclusionIn this study, experiments and inspections were performed to correlate mechanical performance of AM tensile samples to defects observed and measured through scanning electron microscopy. Computed tomography was used for bulk inspection of components and determination of whole-part defect metric distributions. SEM inspections and image analysis tools were used to measure morphology and location metrics of defects present on the fracture surfaces of samples at resolutions higher than that of CT. From the correlation of mechanical performance and defect metrics, the following conclusions can be made.1. Mechanical properties of AM 316L stainless steel are influenced by fluctuations in global energy density varied by laser power.

Mechanical properties of AM 316L stainless steel are influenced by fluctuations in global energy density varied by laser power. Non-optimal energy input density results in the deterioration of mechanical performance as measured by ductility and yield strength. 2. Higher-than-prescribed energy input density is more detrimental to AM 316L stainless steel part performance than is lower-than-prescribed energy input density at the levels tested due to keyhole porosity formed by high energy input.3. When only considering the largest pores observed on a fracture surface, all relationships between pore morphology and mechanical properties found in this study were preserved, indicating the dominating effect of the largest pores on part performance.This work establishes a direct relationship between large, keyhole porosity and decreased mechanical performance, however further investigation is necessary to thoroughly understand this relationship.

Further work to understand the defect-property-performance relationship includes determining how the pore-to-stressed area ratio determines the effect of pores on global part performance, how the conclusions drawn from this study apply to more complex geometries and other materials, how microstructure affects performance, and establishment of minimum critical defect size thresholds for inspection of AM components. These studies will serve to understand the influence of defects on AM part performance and improve the usefulness and safety of additively manufactured components.AcknowledgementsSandia National Laboratories is a multi-mission laboratory managed and operated by National Technology & Engineering Solutions of Sandia, LLC, a wholly owned subsidiary of Honeywell International, Inc., for the U.S. Department of Energy’s National Nuclear Security Administration under contract DE-NA0003525.This paper describes objective technical results and analysis.

e Effects of Specimen Dimensions on the Mechanical Behavior of EBM Produced Ti6Al4V Alloys R.S. Kircher, A.M. Christensen, K.W. Wurth Medical Modeling Inc., Golden, CO 80401AbstractThere are several published studies investigating the microstructures and mechanical properties obtained during additive manufacturing of Ti6Al4V alloys utilizing the Electron Beam Melting (EBM) technique. These studies have concentrated on conventional testing coupon sizes and configurations which allowed for a direct comparison to the properties of conventionally produced Ti6Al4V alloys. One of the many benefits of the EBM process is that it allows the manufacturer to produce components in sizes and configurations unachievable by conventional methods. It becomes important to understand and verify the microstructures and mechanical performance of these smaller components in a manufacturing environment, requiring the use of non‐conventional testing configurations.

This paper presents case-studies involving the production and testing of non-standard samples and how these samples compare to conventional E8 testing coupons. Differences in mechanical performance were observed and are most likely due to the unique characteristics of EBM produced materials.IntroductionThe rapid prototyping industry is in the midst of transition from an industry that produced models, prototypes, and other non-functional objects to an industry that is not only making functional components, but components that will be expected to meet some of the strictest quality standards encountered in the manufacturing industry. The ability to additively manufacture fully melted metal components out of high-end engineering alloys such as Titanium and other super alloys is a powerful tool, but these high-end alloys are typically only used in extreme environments and for critical applications; namely in the biomedical and aerospace industries.

At the very least, a manufacturer must provide its customers with certifications that the additively manufactured components meet a predetermined set of mechanical, chemical and performance properties. In addition to certifications many customers require a thorough quality control system be established, documented and regularly verified to provide a high assurance that the material is consistently meeting the necessary performance criteria.Certifying a batch of conventionally produced material is a relatively straight-forward and well established practice in the manufacturing industry.

For example, if one is going to use wrought Ti6Al4V alloy bar stock to machine a batch of medical implants, they would order a lot of material to a specific standard, such as an ASTM, ISO or AMS standard and the material would be delivered to the manufacturer with the proper paperwork certifying that the proper analyses have been conducted, in the proper manner, and that the results satisfy the necessary values outlined in the standard. Typically these specifications will outline sample frequencies, quantities, configurations, and geometries as well as outline testing conditions and acceptable testing methods. For example, if a lot of wrought ASTM F 1472 Ti6Al4V bar is to be certified, a minimum of one (1) representative chemical analysis must be conducted in addition a tensile test, and a metallographic evaluation [1].

If these evaluations are carried out correctly and the results meet the requirements of the standard, the entire ‘lot’ of material is considered certified for use to that standard. The ASTM F 1472 specification defines a “lot” of material as the total number of mill products produced from one heat under the same conditions at essentially thesame time [1], which means that a single material certification may be encompass a large quantity of raw material, perhaps enough material to allow the production of several hundred medical components.The additive manufacturing industry has yet to adopt a set of terminology, test methods or material specifications to standardize the quality control methods to be used in fabricating and certifying layer-based materials. As a result, the specifics surrounding material certification have typically been negotiated between suppliers and customers on an individual case basis.

As a result, the specifics surrounding material certification have typically been negotiated between suppliers and customers on an individual case basis. Standardizing the certification specifics of additive manufacturing technologies will provide a higher comfort level when using additively manufactured materials in addition to eliminating the need to negotiate material certifications on a single contract basis.The ultimate goal of a material certification is to produce strong evidence that all of the material supplied to a customer meets a specific set of criteria, while minimizing the time and cost of testing necessary to provide that evidence. Additive manufacturing technologies have opened doors by providing a method of producing components with geometries that are unachievable by conventional manufacturing processes. However, this poses an interesting dilemma.

However, this poses an interesting dilemma. Since these are novel part geometries, testing methods to verify the material properties of these components do not exist. In addition to developing and optimizing the manufacturing process necessary to produce a layer-based product, the additive manufacturer must also develop the test methods and acceptance criteria to be used for material certification of that product.This paper will present case studies investigating the effects of tensile specimen geometries on the measured mechanical properties of one particular metal alloy, Ti6Al4V, produced by the Electron Beam Melting process.The Electron Beam Melting ProcessLike many rapid manufacturing techniques, the EBM process creates a physical component from digital CAD models by building the component in a series of layers. The EBM process starts by distributing a 100 µm layer of fine metal powder on a steel platform.

The EBM process starts by distributing a 100 µm layer of fine metal powder on a steel platform. An electron beam is produced by passing current through a Tungsten filament. The electron beam scans areas of the metal powder layer, in an x-y coordinate system, as defined by the computer model fully melting the powder in the areas scanned. Once the beam has scanned the appropriate areas the steel platform is lowered by 100 µm, and a new layer of powder is distributed on top of the previously melted layer. This process continues, layer by layer, until a complete part is produced. During processing the entire build chamber maintains a temperature of approximately 500oC.

This process continues, layer by layer, until a complete part is produced. During processing the entire build chamber maintains a temperature of approximately 500oC. Once the part has been completed, the build chamber is flooded with He gas to expedite cooling.The use of an electron beam to supply the energy necessary to melt the metal powder mandates that the process be done in a vacuum chamber, which minimizes chemical reactions between the melting metal powder and the surrounding atmosphere. This feature is extremely beneficial in producing objects out of Ti6Al4V-ELI material, because the low levels of interstitial elements such as O, N and H can be controlled during production. Currently the EBM process is capable of producing parts up to 200x200x200 mm, with a dimensional accuracy of 0.4 mm.Hot isostatic pressing (HIP) has become an industry standard for investment cast Ti6Al4V alloy components used in the orthopaedic industry.

Since EBM produced components will most likely be used in similar applications, unless otherwise noted, the data shown for test specimens in this study was acquired following standard HIP treatment. A standard HIPtreatment cycle for titanium alloy consisting of two (2) hours at 900oC and 103.4 MPa was utilized [2].Case Study 1: Standard E-8 Testing specimensTo date, the vast majority of tensile data available for additively manufactured metals has been generated using the well established ASTM E-8 standard: Test Methods of Tension Testing of Metallic Materials [3]. Figure 1 is a schematic of a build configuration developed to produce a series of standard E-8 round samples. The vertical cylinders were designed to accommodate tensile testing in the axis parallel to the beam direction (Z-orientation). The large slabs of material were sliced into coupons and machined to accommodate tensile testing in the axis perpendicular to the beam direction (XY-orientation).

The large slabs of material were sliced into coupons and machined to accommodate tensile testing in the axis perpendicular to the beam direction (XY-orientation). The sample size chosen for the analyses were 1” (25.4mm) gauge length round specimens with the dimensions listed under specimen 3 as shown in figure 2 [3].Figure 1: Schematic illustrating the build configuration used to produce standard ASTM E-8 specimens.Table A summarizes the tensile properties obtained from the analysis of a set of test specimens built in the configuration shown in figure 1. Previous studies have shown that chemical composition, especially oxygen and iron content have a significant effect on tensile properties of EBM produced Ti6Al4V material [4].

Previous studies have shown that chemical composition, especially oxygen and iron content have a significant effect on tensile properties of EBM produced Ti6Al4V material [4]. It should be noted that the results in table A were obtained from a powder containing 0.13 wt.% Oxygen and 0.07 wt.% Fe; all other elements were within the required chemical composition requirements of medical grade Ti6Al4V alloy.Figure 2: ASTM Standard E-8 round specimen dimensions. Specimen 3 samples were evaluated in this study.Figure 3 compares the results obtained from these standard E-8 specimens to the standard mechanical property requirements of conventionally manufactured cast and wrought Ti6Al4V products.

These results, as well as results from other studies, suggest that EBM produced Ti6Al4V material exhibits mechanical properties that are in line with those required by the established standards [5].Table A: Summary of the mechanical property data obtained by testing standard E-8 roundspecimens.IDOrientation (XY or Z)Avg. YS (ksi)Avg.

YS (ksi)Avg. UTS (ksi)Elongation (%)RA (%)1XY11613017.055.22Z11613017.747.0140120EBM Solidified & Heat treated ASTM F 1108100ASTM F 13680ASTM B 348 Grade 2360ASTM F 147240200Yield Strength (ksi)UTS (ksi)Reduction of Area (%)Elongation (%)Figure 3: Comparison of EBM produced Ti6Al4V material and conventional Ti6Al4V Standardproperties.Case Study 2: Un-machined Thin Round Specimens Comparing results from standard ASTM E-8 to the properties of other conventionally produced materials is useful as a benchmarking exercise, and standard E-8 specimens may be sufficient to certify the mechanical properties of certain EBM produced Ti6Al4V products. However, the use of standard samples poses a few problems for the additive manufacturer. Generally, the quantity of material produced by an additive manufacturer is magnitudes less than the amount of material produced by a major casting house or rolling mill.

Generally, the quantity of material produced by an additive manufacturer is magnitudes less than the amount of material produced by a major casting house or rolling mill. The cost of acquiring machining and tensile testing capabilities to accommodate the ASTM E-8 test methodology is significant to small-scale production. One of the major advantages of additive manufacturing is that one can design a component and literally have a fully functional part in their hands in the span of a few days. If the testing necessary to certify the properties of that component is not conducted at the manufacturing facility, but rather by an independent testing lab, the time involved with obtaining the necessary data adds a significant, undesirable delay in the manufacturing lead-time.More importantly, standard ASTM E-8 specimens may not be representative of the mechanical properties of EBM produced components with complex geometries, especially components that are smaller than a standard specimen.

One of the most promising applications of additive metal manufacturing is the production of truss-like, lattice structures, such as the component shown in figure 4. These structures allow an engineer to manipulate properties such as the weight, density, and elastic modulus of designed component by changing the configuration of the lattice. The thickness of structural members, or struts, is generally much less than the thickness of a standard testing coupon. The ideal tensile testing coupon for material certification of EBM produced materials would be a coupon that requires little or no machining, and would mimic the size and production environment of actual components.Figure 4: An example of a truss-like structure produced by electron beam melting.The goal of this case study was to investigate the possibility of producing thin, round tensile samples that would be representative of struts in a designed lattice structure.

A secondary goal of the study was to investigate the feasibility of using as-produced Ti6Al4V testing coupons, without any machining prior to tensile testing. To accomplish these goals a series of thin round test specimens, illustrated in figure 5, with varying gauge thicknesses were designed and produced in Ti6Al4V alloy using the EBM process. Test coupons were built in three orientations; Z-orientation, XY orientation, and diagonally (45o angle from beam direction).Figure 5: Schematic showing the design of thin round tensile specimens used in case study 2.The designed sample thickness, EBM produced thickness, and mechanical properties for the three sample sets are presented in Table B. The mechanical properties listed in the table were calculated based on the average measured gauge thickness of the un-machined sample. Comparing the measured sample thickness to the designed sample thickness indicates that the EBM process produced samples thicker than the designed geometries.

Comparing the measured sample thickness to the designed sample thickness indicates that the EBM process produced samples thicker than the designed geometries. Previous studies haveshown that EBM produced components have a inherent surface roughness, and are typically oversized by as much as 0.5 mm on any given surface, especially in the XY plane[4]. What is not well understood is how this additional material influences the mechanical properties of EBM produced components. The effect of this additional thickness and surface roughness on the measured mechanical properties of these specimens is best understood by examining the vertically orientated sample set. Figure 6 plots the measured yield strength and UTS values for the vertically orientated test specimens as a function of designed sample thickness. In addition to the measured strengths, a theoretical strength is plotted.

In addition to the measured strengths, a theoretical strength is plotted. The theoretical strength was determined by using the designed sample thickness in lieu of the measured sample thickness when calculating the cross-sectional area of the gauge section.Some interesting conclusions can be made by comparing these two strength values. Consider the case where the additional gauge thickness measured in these samples consisted of a uniform layer of fully dense, smooth material with the same properties as the rest of the material. If this were the case, the measured strength values would not be influenced by the thickness of the designed gauge section. Since the measured strength drops considerably as the designed sample thickness is decreased, it can be concluded that this additional surface material does not possess the same load-bearing capabilities as the base material. Now consider the case where the additional gauge material offers no load-bearing capacity.

Now consider the case where the additional gauge material offers no load-bearing capacity. If this were the case, the theoretical strength values would not be affected by the designed gauge thickness. This is obviously not the case, as figure 6 indicates that the thin gauge specimens are capable of bearing much more load than expected. The implication of these results is that if an engineer were to design an EBM produced lattice structure consisting of 1mm struts, the structure would actually support more load than the design would indicate. This is because the struts in the as-produced component would be larger than 1mm when produced by the EBM system. However, if a material testing lab was presented with the same EBM produced lattice component, and followed conventional methods for determining strength values they would conclude that the material properties of the component were below the expected values for Ti6Al4V alloys.

The sample geometry used to certify a given product is dependent on how the material is produced, and the final dimensions of product. Thin sheet steels are certified using thin flat ‘dog-bone’ sample geometries, while large Ti6Al4V bar stock is certified using large round tensile specimens. Additive manufacturing poses an interesting dilemma with respect to the selection of sample geometries. The testing coupons produced for material certification of additively manufactured components are typically produced separately allowing for virtually and test sample configuration. The goal of this case study was to compare the mechanical properties of EBM produced Ti6Al4Vmaterial obtained using standard E-8 flat ‘dog-bone’ testing specimens (figure 7[2]) to results obtained using the standard 1” gauge length round samples. The necessary samples were harvested out of large blocks of EBM produced Ti6Al4V material and were tested along in the XY Orientation.

The necessary samples were harvested out of large blocks of EBM produced Ti6Al4V material and were tested along in the XY Orientation. The material tested in this study was not subjected to HIP prior to testing.Table C summarizes the results obtained from during this study. The sample size in this study was not sufficient to draw any strong conclusions, as the values for both testing specimen geometries are rather close, but the round specimens are slightly stronger. A larger scale study will be necessary to determine if a difference in the results obtained between the sample geometries exists.Figure 7: Flat E-8 specimen geometryTable C: Summary of the mechanical properties obtained from both flat and round standard ASTM E-8 specimensSample ConfigurationNumber of SamplesAvg. YS (ksi)Avg. UTS (ksi)Reduction of Area (%)Avg. Elong.

YS (ksi)Avg. UTS (ksi)Reduction of Area (%)Avg. Elong. (%)Flat61201343315Round61231353917ConclusionsThis paper presents case studies that investigate the effects of sample geometries on the measured mechanical properties of EBM produced Ti6Al4V alloys. Utilizing standard round ASTM E-8 test specimens it was shown that EBM produced materials have properties similar to conventionally produced Ti6Al4V alloys.Although standard E-8 specimens provide useful information about the overall capabilities of the EBM production process, they provide some logistical challenges for the additive metal manufacturer, and they do not necessarily represent the mechanical behavior one can expect from EBM produced components with complex geometries. Particularly, truss-like structures pose an interesting dilemma as they are much thinner than conventionally produced test specimens, and have an inherent excess surface roughness.

Particularly, truss-like structures pose an interesting dilemma as they are much thinner than conventionally produced test specimens, and have an inherent excess surface roughness. This excess surface roughness makes it difficult to predict and test mechanical performance of a lattice structure because it increases the cross-sectional area of the structural members, but the extent of its load-bearing contribution is unknown. The use of flat specimens opposed to round specimens may yield slightly lower yield and UTS; however additional testing will be necessary to show a statistically significant difference.The ultimate goal of the additive manufacturing industry should be to develop a standard testing methodology tailored to unique challenges and capabilities associated with layer based deposition of material. These case studies provide information that can be used to aid further studies in developing such standards.1. Annual Book of ASTM Standards, Vol.

viewed PaperEFFECTS OF SPHERICAL FILLERS ON THE PROCESSABILITY AND MECHANICAL PROPERTIES OF PA613 AND PP-BASED LASER SINTERING DRY BLENDSI. Kletetzka*, R. Gawlikowicz*, H.-J. Schmid*Direct Manufacturing Research Center (DMRC) and Particle Technology Group, Paderborn University, GermanyAbstractPolymer materials filled with particles may show substantially altered mechanical properties. Therefore, it is an important aim to be able to tailor the mechanical properties of LS components by adding fillers and thus to create new application areas for additively manufactured components. In this work, the influences of spherical fillers on the processing properties and the resulting mechanical properties of laser-sintered components are investigated. For this purpose, micro glass spheres, hollow glass bubbles and mineral spheres are dry blended to the matrix polymers polyamide 613 and polypropylene with filling ratios of 20 and 40 vol%.

For this purpose, micro glass spheres, hollow glass bubbles and mineral spheres are dry blended to the matrix polymers polyamide 613 and polypropylene with filling ratios of 20 and 40 vol%. First, relevant properties of the blends, such as powder flowability, thermal behavior and melt viscosity, are investigated. Based on the results, processing parameters are then developed for the laser sintering (LS) process and the mechanical properties of the components are investigated.IntroductionAdditively manufactured components are increasingly used as functional prototypes or as end user products [1, 2]. In the field of plastics, selective laser sintering appears to be particularly interesting for this purpose, as it enables the production of complex geometries without the use of support while offering a comparatively high productivity [3]. In selective laser sintering, thermoplastic powder is deposited in layers, preheated and selectively melted by a laser to form plastic components.

In selective laser sintering, thermoplastic powder is deposited in layers, preheated and selectively melted by a laser to form plastic components. For producing end user products, functional and mechanical properties of additively manufactured components are becoming more important. However, currently there is only a limited selection of LS materials available, meaning that not all customer-specific requirements can be met [2]. Particularly in the automotive and aerospace industries, filled plastics represent standard materials, as they can exhibit better mechanical properties, higher heat resistance or improved wear properties, depending on the filler [4, 5].In order to create new fields of application for selective laser sintering, the material variety should be extended by filled variants (dry blends) of polypropylene (PP) and polyamide 613 (PA613) within this work.

For this purpose, the influence of different fillers and the filling ratio on the mechanical properties will be analyzed, thus enabling application specific tailoring of material properties. The main goal here is an increase in component stiffness while maintaining reasonably good elongation at break and tensile strength. The aim is to not only increase the component stiffness in the dry state, but also after conditioning the test specimens.219The term filler covers a wide range of materials. Fillers are defined by various authors as natural or synthetic particles that do not dissolve in the base polymer and are not volatile [5–7]. Consequently, a matrix is formed of polymer and disperse filler particles with distinct interfaces. The particle shape of the filler can be either spherical, acicular, flakey, fibrous or irregular. At current state of the art, there are essentially three different methods for producing filled powders for the LS process.

At current state of the art, there are essentially three different methods for producing filled powders for the LS process. These methods differ significantly in the properties of the powders produced and in the effort required to produce the respective powders. So far, there are only relatively sparse data available comparing the powders and the resulting component properties of the different manufacturing methods and regarding the impact of different fillers in general. [4, 8–15]The easiest method for creating filled SLS powders is dry blending. Polymer particles and filler particles are mechanically mixed, for example in a drum hoop mixer or a rotary mixer. Different authors used mixing times in the range of 30 min up to 24 h [10, 13]. With dry blending, a commercially available plastic powder can be quickly mixed with different fillers and filling ratios, allowing the material to be adapted to a specific build job.

With dry blending, a commercially available plastic powder can be quickly mixed with different fillers and filling ratios, allowing the material to be adapted to a specific build job. A possible disadvantage of dry blending is an inhomogeneous distribution of the filler in the polymer powder. Especially if polymer and filler particles differ strongly in morphology and density, the preparation of a homogeneous blend and preventing segregation upon processing is challenging. In addition, the adhesion of filler particles and polymer matrix is based solely on pressure less coalescence in the build process. [15]An alternative method is the production of filled SLS powder by melt compounding and subsequent cryogenic milling to form a powder. The achievable particle size distribution is often wide and sieving or other classification processes are necessary after milling.

The achievable particle size distribution is often wide and sieving or other classification processes are necessary after milling. Furthermore, the particles tend to have rough surfaces and a very irregular morphology leading to a poor powder flowability. An additional rounding process or the addition of flow improving additives might be necessary. Another challenge is that during the grinding process, separation of the polymer and filler, or comminution of the filler, can take place. [3, 11, 14, 16, 17]The production of a filled LS powder via thermally induced phase separation (TIPS) aims to create particles in which the filler is completely coated with plastic. Advantages are, that the generated powder usually exhibits a good flowability. However, the process is complex and requires precise control of the process parameters to achieve the desired powder properties.

However, the process is complex and requires precise control of the process parameters to achieve the desired powder properties. Furthermore, the use of solvents at high temperatures can only be realized with considerable effort in terms of occupational safety and environmental protection. [8, 18]MaterialsThe intended approach is to produce a filled material by dry blending various combinations of commercially available polymer powders and fillers. Consequently, the materials must already be available in powder form with an appropriate particle size distribution. Polypropylene (PP 01 nat by BASF) and Polyamide 613 (VESTOSINT 3D 8754HT1 by Evonik) are selected as base polymers for the dry blends. For these polymers, the properties of dry blends have not been studied systematically. Furthermore, PA613 offers better mechanical properties without a filler when compared to PA12.

For these polymers, the properties of dry blends have not been studied systematically. Furthermore, PA613 offers better mechanical properties without a filler when compared to PA12. PP has the advantage, that it absorbs almost no moisture and the effects of moisture exposure on mechanical properties should therefore be minimal [19, 20].220This work focuses on spherical fillers, since fibrous fillers proved to cause high anisotropy when used in dry blends for laser sintering and can furthermore decrease the powder flowability [21–23]. Three different spherical fillers with suitable particle size distributions were chosen: glass beads manufactured by Sigmund Lindner (Silibeads 40-70 [24]), glass bubbles manufactured by 3M (iM16k [25]) and mineral beads manufactured by Itochu Ceratech (Naigai Cerabeads #1700 [26]). The Silibeads are made of recycled soda lime glass (A-glass) and have a density of 2.5 g/cm³.

The Silibeads are made of recycled soda lime glass (A-glass) and have a density of 2.5 g/cm³. 3M glass bubbles are also made from soda lime glass but due to being hollow, their apparent density is only 0,46 g/cm3. The isostatic crush strength of iM16K is 110 MPa. Naigai Cerabeads are made of sintered mullite and have a density of 2,9 g/cm3. Both glass fillers were ordered with a silane coating recommended by the manufacturer for the used polymers and without coating. For Silibeads a gamma-aminopropyltriethoxysilane coating with the designation Si-3 was selected, for iM16K the coating was not specified by the manufacturer.The polymer powders and fillers were mixed using a drum hoop mixer. Each dry blend was mixed for 30 min at a rotational speed of 30 rpm. Since the selected fillers differ considerably in terms of density, the filler content is given in percent by volume. This allows the best possible comparability of the results between the individual filler types.

This allows the best possible comparability of the results between the individual filler types. PA613 und PP are blended with the different fillers (with and without silane coatings) in filling ratios of 20 vol% and 40 vol% as listed in table 1. Additionally, all tests are carried out with the unfilled polymers for comparison.Table 1: Overview of the investigated dry blendsPolyamide 613 (VESTOSINT 3D 8754HT1)Polypropylene (Ultrasint PP 01 nat)Glass beads (Silibeads) uncoated and silane coated0 / 20 / 40 vol%0 / 20 / 40 vol%Glass bubbles (iM16K) uncoated and silane coated0 / 20 / 40 vol%0 / 20 / 40 vol%Mullite beads (Cerabeads) uncoated0 / 20 / 40 vol%0 / 20 / 40 vol%Experimental MethodsBefore the effect of the fillers on the mechanical properties of laser-sintered components can be identified, suitable parameters must first be developed in order to process the respective blends.

For this purpose, the methodology shown in figure 1 was used.Preliminary tests are carried out on the powder to characterize the impact of different fillers on the powder properties. Subsequently the parameter development on the SLS Systems was carried out. Here, the parameter sets for unfilled polymer powders were used as a starting point. The specimens built with varying laser energy are used to analyze the mechanical properties, part resolution and density. The results are used to determine a feasible laser energy density.221Figure 1: Methodology for parameter development for dry blendsFor studying the particle morphology, a Phenom World scanning electron microscope (SEM) was used, which allows magnification up to 24,000 times. Beforehand, the samples were sputtered to create the electrical conductivity required for SEM imaging.The particle size distribution was analyzed using dynamic image analysis for moving particles according to ISO 13322-2.

The measurements were carried out with a QICPIC manufactured by Sympatec, equipped with a dry dispersion unit. The set measuring range was M5, allowing the detection of particles with a size in between 1,8 and 3755 μm. [27]The flowability of the powder and mixtures was evaluated by means of the Hausner ratio HR. The description of the Hausner ratio is given in VDI 3405 Sheet 1.1. The ratio is defined as the quotient of the tapped density and the bulk density. For measuring, a cylinder was filled up to the 250 ml mark with the powder mixture to be tested, before the powder was deagglomerated in a plastic bag. The measuring cylinder was then closed and tapped uniformly with low force onto a solid surface at a frequency of 0.5 Hz for a duration of two minutes. The process is then repeated with a duration of one minute until the tapping volume no longer changed.

The process is then repeated with a duration of one minute until the tapping volume no longer changed. To improve the measurement accuracy, the Hausner factor was determined three times for each data point and the mean value was calculated. [28]The thermal behavior of the dry blends was investigated by differential scanning calorimetry (DSC) according to DIN EN ISO 11357-3. A DSC 214 Polyma from Netzsch was used as the measuring instrument. In DSC analysis, small sample quantities (5±0.3 mg) are heated up to a defined temperature and cooled down again, whereby the heat flow between the sample and an empty reference crucible is continuously measured and recorded. During the investigations, the samples were heated up to 250 °C at a heating and cooling rate of 10 K/min. [29]For the determination of the melt volume-flow rate (MVR), a capillary rheometer of the manufacturer Zwick Roell is used. The test procedure is described in DIN EN ISO 1133-1.

[29]For the determination of the melt volume-flow rate (MVR), a capillary rheometer of the manufacturer Zwick Roell is used. The test procedure is described in DIN EN ISO 1133-1. Thus, the material to be tested is filled into a cylinder which is heated and melted with a duration of 5 minutes. A piston loaded with a weight generates a pressure which forces the melt through an extrusion nozzle. Before testing, the powder samples were dried in a laboratory oven for 30 min at 105°C to avoid bubble formation and hydrolysis. The test temperature for PA613 powder blends was 240 °C. For PP powder mixtures, the test temperature was set to 220 °C. For all MVR measurements, a weight of 5 kg was used. For each blend, a minimum of three tests was carried out and the average MVR value was calculated. [30]222For the density measurement of the sintered components, cubes measuring 20 x 20 x 20 mm were built and tested using the Archimedes principle, which is standardized by DIN EN ISO 1183-1.

[30]222For the density measurement of the sintered components, cubes measuring 20 x 20 x 20 mm were built and tested using the Archimedes principle, which is standardized by DIN EN ISO 1183-1. A sample attached to a wire is first weighed in air and then in ethanol. For this purpose, a precision scale from Sartorius with the model name CPA224S and the associated set for density determination with the name YDK01 was used. For each parameter set, the average density was calculated using three test specimens. [31, 32]Tensile tests according to DIN 527 were performed to determine the mechanical properties such as modulus of elasticity, tensile strength and elongation at break. For this purpose, laser sintered specimens of type 1A and universal testing machines of type Instron 5569 and Zwick Roell Z2020 were used. The test speed for determining the modulus of elasticity is 1 mm/min and is then increased to 50 mm/min at an elongation at 0.25 % until fracture.

The test speed for determining the modulus of elasticity is 1 mm/min and is then increased to 50 mm/min at an elongation at 0.25 % until fracture. In contradiction to the material standard for polyamide and polypropylene, both unreinforced and filled materials are tested at 50 mm/min to ensure the best possible comparability. For every parameter set, average mechanical properties were calculated using 5 test specimens. [33–35]The mechanical properties of polyamides are strongly dependent on the moisture content in the material, therefore a precise adjustment of this value is necessary to obtain comparable measurements. Tensile properties were therefore evaluated in the dry state and after conditioning. Due to time constraints, accelerated conditioning according to DIN EN ISO 1110 was carried out in a climatic chamber at 70 °C and 62 % relative humidity.

Due to time constraints, accelerated conditioning according to DIN EN ISO 1110 was carried out in a climatic chamber at 70 °C and 62 % relative humidity. To prevent the absorption of moisture in the dry state specimens, they were kept dry in a plastic bag with silica gel pads after the build job until the tensile test was performed. [36]Powder CharacterizationThe size, shape and the surface of powder particles have a decisive influence on the powder behavior during the SLS process. If the powder particles have an irregular shape or rough surfaces, the flowability of the powder decreases, since the adhesive forces between the individual particles increase. At the same time, the packing density of the powder decreases. Poorer powder flowability can lead to uneven coating during the powder application process, and reduced powder bed density can lead to pores and a decrease in the mechanical strength.

Poorer powder flowability can lead to uneven coating during the powder application process, and reduced powder bed density can lead to pores and a decrease in the mechanical strength. Moreover, the particle size distribution has an influence on the surface quality and the detail resolution of the final part. A high proportion of large powder particles has a negative effect on both properties. An excessive fine powder content, however, has a negative effect on flowability and pourability [3, 37].The first filler considered is uncoated glass beads named Silibeads. The SEM picture shows, that Silibeads have a very smooth particle surface (figure 2). Most of the particles have a spherical shape, but some irregular particles are present. In addition, some small particles are partially molten into the surface of larger particles. The average sphericity of the particles is 0.852. Silibeads have a relatively narrowly distributed particle size between about 25 µm and 100 µm.

The average sphericity of the particles is 0.852. Silibeads have a relatively narrowly distributed particle size between about 25 µm and 100 µm. In terms of particle morphology and particle size distribution, the Silibeads should be suitable as a filler for the laser sintering polymer powders. Silibeads with a silane coating showed no obvious differences in terms of particle size distribution and particle morphology.2234PSD Silibead uncoated3,532,5)₃ q ( gl21,510,500255075 xm / µm100125150Figure 2: Silibeads uncoated. SEM Picture on the left and particle size distribution on the rightPSD Cerabeads#1700 uncoated54)3₃ q ( gl2100255075 100 xm / µm125150Figure 3: Cerabeads#1700. SEM Picture on the left and particle size distribution on the rightPSD iM16K uncoated32,52)₃ q ( gl1,510,500255075 xm / µm100125150Figure 4: iM16K uncoated.

SEM Picture on the left and particle size distribution on the rightPSD iM16K uncoated32,52)₃ q ( gl1,510,500255075 xm / µm100125150Figure 4: iM16K uncoated. SEM Picture on the left and particle size distribution on the rightThe surface of the Cerabeads is significantly rougher compared to the Silibeads, as can be seen in the SEM image in figure 3. However, the surface is free of cracks. The particle morphology can be classified as spherical to slightly "potato-shaped". This is reflected in a value for the average sphericity of 0.823. Cerabeads have a larger average particle size compared to Silibeads. The D90224175175175value of Cerabeads is 120.13 µm, which means that 10 % of the particles have a diameter which is bigger than the typical layer height of 120 µm.

The D90224175175175value of Cerabeads is 120.13 µm, which means that 10 % of the particles have a diameter which is bigger than the typical layer height of 120 µm. The high proportion of coarse particles potentially increases the surface roughness of laser-sintered components and could be problematic during processing.The SEM images of the hollow glass spheres (iM16K uncoated) already show that they are much smaller than the other fillers (fig. 4). The D50 is 18.4 µm. For the most part, the glass spheres have an almost perfectly spherical shape. Under the scanning electron microscope, some broken glass bubbles can be seen. These have an uneven, flakey morphology. Smaller particles adhere to the surface of the hollow glass spheres, which can be seen well under higher magnification. The mean sphericity measured in the QICPIC is 0.829.

Smaller particles adhere to the surface of the hollow glass spheres, which can be seen well under higher magnification. The mean sphericity measured in the QICPIC is 0.829. This value is at disagreement with the spherical particle shape that can be seen by SEM and is possibly caused by the percentage of fractured glass. Due to the small particle size the powder flowability could be degraded by the addition of iM16K. Glass bubbles with a silane coating showed no obvious differences in terms of particle size distribution and particle morphology.In addition to the particle morphology and the particle size distribution, the flowability of the powder is another important parameter to ensure a good powder layer application and a high powder bed density in the LS process. As already described, there are correlations between flowability, particle size distribution and particle morphology.

As already described, there are correlations between flowability, particle size distribution and particle morphology. In the scope of the investigations, the Hausner ratio was used to describe the flowability of powders and blends. If the Hausner ratio is below 1.25, the powder is supposed to have good flowability. With a ratio above 1.4, cohesion dominates and it is assumed that the powder tends to be unsuitable for the SLS process. Between a ratio of 1.25 and 1.4, the powder is classified as having limited flowability. [28]The results for the powder flowability of dry blends in dependence on the filling ratio is given by figure 5. Starting with the results for PP based dry blends, it can be noted, that the addition of a filler materials with a filling ratio of 20 vol% improves the powder flowability for all tested fillers. Even hollow glass spheres (iM16K) improve the flowability compared to unfilled PP although they exhibited a worse flowability when tested separately.

Even hollow glass spheres (iM16K) improve the flowability compared to unfilled PP although they exhibited a worse flowability when tested separately. In contrast, the addition of 40 vol% of uncoated iM16K worsens the flowability. The greatest improvements can be achieved with the addition of 40 vol% Cerabeads or silane coated Silibeads. These fillers also showed the best flowability when measured separately. The Hausner ratio of the blends is below 1.25 in both cases, therefore the powder achieves a good flowability in contrast to the unfilled PP powder.The results for PA613 based dry blends show the same trends as the results for PP (fig. 5). However, unfilled Vestosint PA613 has a better flowability when compared to PP. The influence of a 20 vol% filling ratio on the flowability is minor. All blends still have a good flowability. For a filling ratio of 40 vol%, the impact on the flowability is more pronounced.

The influence of a 20 vol% filling ratio on the flowability is minor. All blends still have a good flowability. For a filling ratio of 40 vol%, the impact on the flowability is more pronounced. Silibeads and Cerabeads can again be used to slightly increase the flowability. Uncoated iM16K worsen the flowability to the point, where the powder is classified as having a reduced flowability. Again, the silane coated fillers tend to perform slightly better.225Flowability of PP based blendsPA6-13 based blends1,361,281,341,261,321,24o i t a rr e n s u a H1,31,281,221,21,181,261,161,241,141,221,120%10%20% Filling ratio30%40%0%10%20%Filling ratio30%40%Figure 5: Hausner ratio of PP and PA613 based dry blendsIn order to investigate the thermal behavior of the polymer powders and the powder blends, the respective temperature-process windows were determined by using DSC analyses.

In addition to the effects on the temperature process window, other effects of the fillers on the heat capacity and the thermal conductivity of the powder bed are possible.Figure 6: Thermal processing windows of PP and PA613 based dry blends with 40 vol% fillerThe temperatures at the start of recrystallization and the start of the melting process have been extracted from the DSC curves and plotted in figure 6. At first glance, it can be seen that a narrowing of the process window takes place for all PA613 powder blends compared to unfilled226SiLiBeads uncoatedSiLiBeads siiM16k siiM16k uncoatedCerabeadsPA613. For all blends the start of the melting process remains unchanged compared to the unfilled material as expected. Only the recrystallization temperature is changed by the addition of fillers. Most likely fillers have a nucleating effect, which causes the PA613 crystals to grow earlier [14]. The process window for unfilled PA613 is around ∆T = 39 °C.

Most likely fillers have a nucleating effect, which causes the PA613 crystals to grow earlier [14]. The process window for unfilled PA613 is around ∆T = 39 °C. The largest impact on the process window is observed for PA613 + 40 vol.% Silibeads uncoated. The processing window of the dry blend is narrowed to ∆T = 23.8 °C. The smallest impact on the processing window is observed for PA613 + Cerabeads 40 vol.%. The process window here has a span of ∆T = 31.1 °C. In the comparison between the silane coated and uncoated Silibeads and iM16k, respectively, the silanization seems to have a slightly positive effect on the size of the temperature process window.In addition to the impact on the processing window a reduction in heat capacity and a reduction of the measured melting and recrystallization enthalpy were observed for the PA613 powder blends.

The area under the peaks of the DSC curve is correspondingly smaller, since the filler does not melt and thus the enthalpy of melting of the mixture is lower. Heat capacity and thermal conductivity of polymer and filler are summarized in table 2.The thermal processing windows for unfilled PP as well as for PP powder blends are also given by figure 6. In contrast to PA613, the fillers have no significant influence on the process window. Accordingly, the fillers do not seem to influence the recrystallization. A comparison of the coated and uncoated fillers also shows no difference. All the processing windows are at about ∆T ≈ 30 °C.

Accordingly, the fillers do not seem to influence the recrystallization. A comparison of the coated and uncoated fillers also shows no difference. All the processing windows are at about ∆T ≈ 30 °C. Similar to PA613, a decrease of the heat capacity as well as a reduction of the area under the peaks in the heating and cooling curves could be observed for the PP powder blends.Table 2: Thermal properties of base polymers and fillersGravimetric Heat capacity [kJ/(kg*K)]Volumetric heat capacity [J/(cm³*K)]Thermal conductivity [W/(m*K)]Melting Enthalpy [kJ/kg]PA613 (Vestosint) [6]1.700.820.23106PP (Ultrasint) [6]2.000.820.2078Glass beads (Silibeads) [38]0.751.091.00Glass bubbles (iM16k) [38]0.750.190.15Mullite (Cerabeads) [26]0.991.611.45In summary, a clear influence of the filler on the thermal process window was found for PA613 blends.

The reduction of the process window can correspondingly worsen the processing properties and cause build job failures due to curling. A new determination of the non-curl temperature is necessary in any case. At first glance, the addition of fillers has no effect on the thermal process window of PP blends. However, the addition of fillers changes the heat capacity and thermal conductivity of the powder. In laser sintering, the heat capacity of the powder in relation to its volume is particularly decisive, since the laser heats a volume element in order to melt it. The required laser power in the process is often specified as laser volume energy density.

The required laser power in the process is often specified as laser volume energy density. The relationship between volumetric 𝑐𝑉𝑜𝑙 and gravimetric heat capacity 𝑐𝑔𝑟𝑎 is given by the powder bed density 𝜌𝑝,𝑏.𝑐𝑉𝑜𝑙 = 𝑐𝑔𝑟𝑎 ∗ 𝜌𝑝,𝑏(1)227The energy required to melt a volume element 𝐸𝑉, neglecting heat losses to other volume elements, is given by the temperature difference before and after Laser interaction ∆T multiplied with the volumetric heat capacity of polymer 𝑐𝑉𝑜𝑙,𝑃 and filler 𝑐𝑉𝑜𝑙,𝐹 and the volumetric melting enthalpy ∆H𝑀,𝑣𝑜𝑙. The filling ratio is accounted for by the factor 𝑉𝑓.𝐸𝑉 = ∆T ∗ [𝑐𝑉𝑜𝑙,𝑃 ∗ (1 − 𝑉𝑓) + 𝑐𝑉𝑜𝑙,𝐹 ∗ 𝑉𝑓] + (1 − 𝑉𝑓) ∗ ∆H𝑀,𝑣𝑜𝑙With the aid of this equation, it is possible to estimate how the required laser energy volume density changes when processing dry blends. Glass hollow spheres, for example, have only about a quarter of the heat capacity of polymer powders and also do not absorb any melting enthalpy.

Glass hollow spheres, for example, have only about a quarter of the heat capacity of polymer powders and also do not absorb any melting enthalpy. It can therefore be estimated, that the required volumetric energy density will be reducing for an increasing filling ratio 𝑉𝑓. Cerabeads on the other hand have almost double the volumetric heat capacity of the polymer powders. This means that the required volumetric energy density can either increase or decrease depending on the temperature difference ∆T. In addition to the model-based differences in the required energy density, the differences in thermal conductivity have additional influences on the melting and cooling behavior of parts and therefore impact the processability.The viscosity of the plastic melt is of great importance for the SLS process, since the molten particles must flow into each other without additional pressure being applied.

Consequently, coalescence is based on surface tension, melt viscosity and gravitational force [3]. In order to characterize the flow behavior, MVR measurements were performed.MVR - filled PPMVR - filled PA613454540403535) n m 0 1 / ³i30253025m c (2020/ R V M1515101055000%10% Filling ratio / vol%20%30%40%0%10%20% Filling ratio / vol%30%40%Figure 7: MVR values for powder blends depending on the filler content; test temperature for PP 220 °C and 240°C for PA613; test weight 5 kg; sample conditioning: 30 min at 105 °CFigure 7 shows the impact of different fillers and filling ratios on the MVR-Value of PP and PA613 based dry blends. As expected, the addition of fillers generally lowers the MVR values since fillers do not melt and thus impair the flowability of the melt. For unfilled PP (mixed powder 60/40), the initial value at 220 °C test temperature and 5 kg test weight is 36.1 cm³/10min.

For unfilled PP (mixed powder 60/40), the initial value at 220 °C test temperature and 5 kg test weight is 36.1 cm³/10min. By adding a volume fraction of 20 % filler, the MVR decreases to 14 – 20 cm³/10min. The MVR value further decreases to 4 – 10 cm³/10min if 40 vol% of filler is added. Cerabeads have the biggest228(2)Silibeads Si3 coatedSilibeads uncoatediM16K Si coatediM16K uncoatedCerabeadsimpact on the MVR value and the glass bubbles (iM16K) have the smallest impact. Furthermore, the silane coated variants of glass beads and glass bubbles tend to impact the MVR value slightly less compared to the uncoated variants. For PA12 powder, it is known that aged powder with a high viscosity tends to form surface defects such as “orange peel” [39]. In particular, for a high filling ratio a negative influence of the high MVR value on the component properties could be possible.

In particular, for a high filling ratio a negative influence of the high MVR value on the component properties could be possible. However, it should be noted here, that the increase in viscosity observed in a nozzle flow for a filled polymer melt is not fully representative for the coalescence process upon laser sintering: This decrease in MVR, i.e. increase in macroscopic viscosity, is obtained while the viscosity of the matrix polymer is basically unchanged. Therefore, it is expected that the deterioration of the coalescence process is less than indicated by the reduced MVR.For unfilled PA613 (virgin powder), the initial MVR value at 220 °C test temperature and 5 kg test weight is 39.6 cm³/10min. As already observed for PP blends, the melt volume flow rate decreases with increasing filler content. Again, glass bubbles (iM16K) had the smallest impact on the MVR values and Cerabeads had the largest impact at a filling ratio of 40 vol% decreasing the MVR value to only 5 cm³/10min.

Again, glass bubbles (iM16K) had the smallest impact on the MVR values and Cerabeads had the largest impact at a filling ratio of 40 vol% decreasing the MVR value to only 5 cm³/10min. The differences between coated and uncoated variants of the same filler mostly fall in the range of the standard deviation for PA613 based blends and are therefore not significant. For both base materials, the influence of the filler on the MVR value decreases with decreasing filler diameter.Parameter Development and ProcessabilityThe preliminary investigations on the powder have shown that the sintering window and the energy required to melt the powder can be changed by the addition of fillers. With regard to the powder flowability, only blends with hollow glass beads showed negative effects. Furthermore, the addition of fillers strongly reduces the MVR value.

With regard to the powder flowability, only blends with hollow glass beads showed negative effects. Furthermore, the addition of fillers strongly reduces the MVR value. The essential parameters which are therefore adapted for the production of test specimens from dry blends are the build chamber temperature and the laser volume energy density.Figure 8: Build job layout for laser parameter development229Since the preliminary powder characterizations showed that the variants with silane coatings performed slightly better regarding relevant powder properties and additionally should provide better mechanical properties, the decision was made to focus on the coated variants. For each blend, initially the same build job was built to find suitable process parameters. The layout of this build job is shown in Figure 8.For each blend, the first step was the determination of the non-curl temperature. For this purpose, the curl cross method was used.

The layout of this build job is shown in Figure 8.For each blend, the first step was the determination of the non-curl temperature. For this purpose, the curl cross method was used. The build chamber temperature for unfilled PP is 138°C and for unfilled PA613 206.5 °C. If curling is detected for multiple curl crosses, the parts are removed, the build chamber temperature is increased by 2°C and a new attempt is made to build curl crosses. If only slight curling occurs for curl crosses in the corners of the build area, the temperature is increased by 1°C and a new set of curl crosses is built until none of the crosses showed visible signs of curling.After determining a feasible build chamber temperature with the curl cross method, different test specimens are built with varying laser energy densities. To limit the number of variables and achieve parameters for a good machine productivity, the laser energy density is only varied by varying the laser power.

To limit the number of variables and achieve parameters for a good machine productivity, the laser energy density is only varied by varying the laser power. The layer height is kept constant at 120 µm and the distance between two hatch lines is kept at 0.25 mm. The scan speed is kept constantly high at 4750 mm/s.The selection of the best laser energy density is based on tensile strength, elongation at break, density, and a subjective assessment of resolution and surface quality. The tensile properties were determined in the dry state without conditioning the test specimens. These are only used to compare the different parameters, so conditioning is not necessary.

The tensile properties were determined in the dry state without conditioning the test specimens. These are only used to compare the different parameters, so conditioning is not necessary. The final mechanical properties will be presented later in figures 11 and 12.PA613 + 40 vol% Silibeads SiDensity604,541,600a P M/t h g n e r t S el i s n e T5040302010Tensile StrengthElongation at Break3,532,521,510,5% /k a e r B t a n o i t a g n o El³m c / g /y t i s n e D1,5801,5601,5401,5201,5001,48002025303540Laser Power / W455001,46024 Laser Power / W3240Figure 9: Mechanical properties and part density of PA613 with 40 vol% glass beads (Silibeads silane coated) in dependence of the laser power23048Figure 9 plots tensile strength, elongation at break and density for different laser powers for a blend of PA613 with 40 vol% Silibeads with silane coating.

It turns out that the mechanical properties of the blend remain almost the same for a relative wide range of laser powers. Looking at the mechanical properties only, laser energies ranging from 24 W to 40 W seem to be feasible for the given scan speed (4750 mm/s), hatch distance (0.25 mm) and layer height (0.12 mm). An optimal part density is achieved for 40 W laser power, however the value for 32 W is almost as good and showed less deviations.The subjective evaluation of the component quality takes into account, among other things, the resolution of the marking, existing surface irritations, surface defects due to layer time jumps, but also the edge and corner quality. These variables were each given a quality rating of 1 to 5. Figure 10 shows the detail resolution and surface quality for PA613 blended with 40 vol% Silibeads Si3 coated. At first it can be noted that the addition of 40 vol% of glass beads gives the part a grayish color.

At first it can be noted that the addition of 40 vol% of glass beads gives the part a grayish color. Regarding the labeling, differences with increasing laser power are clearly visible. The letters and numbers are more precise at low laser power and degrade for the parameter set with higher laser power until they are barely recognizable at a laser power of 48 W. Furthermore, with increasing laser power, the effects of layer time jumps become visible, which negatively affects the resolution of the labeling (z-direction) of the density measurement cubes. The density cubes are particularly prone to surface defects in z direction since they have a relatively high volume and therefore cause a high local energy input. Additionally, with increasing laser power, a decrease in shape and dimensional accuracy with respect to corners and edges can be observed.

Additionally, with increasing laser power, a decrease in shape and dimensional accuracy with respect to corners and edges can be observed. In terms of component resolution, the parameters with a laser power of 24 W and 32 W were rated best.Figure 10: Part resolution and surface quality of PA613 with 40 vol% glass beads (Silibeads silane coated) in dependence of laser the powerTaking into account all the results (mechanical properties, density, resolution), it is found that the optimum laser energy for this blend is around 32 W. However, the results can be further optimized by varying other laser parameters and optimizing the contour exposure separately. Feasible exposure parameters were developed for all different dry blends using the same method.231Table 3: Processing parameters of PA613 based blendsPA613unfilled20 vol% Silibeads40 vol% Silibeads20 vol% iM16k40 vol% iM16k20 vol% Cerabead40 vol% CerabeadBed temp.

[°C]206.5206.5206.5207208206.5209Laser energy [W]38323224243232Table 4: Processing parameters of PP based blendsPPunfilled20 vol% Silibeads40 vol% Silibeads20 vol% iM16k40 vol% iM16k20 vol% Cerabead40 vol% CerabeadBed temp. [°C]138.5138.5138.5140142139.7Laser energy [W]27.5222216,516.537Mechanical PropertiesThe results for the PA613 based blends built with optimized parameters as described above are displayed in figure 11. First of all, it can be seen that with all fillers used, the stiffness in the dry state could be increased. With increasing filler content, the stiffness also increases. The highest increase can be achieved with the addition of 40 vol% Cerabeads. Here, the stiffness could be increased from 2162 MPa to 7550 MPa. The smallest increase in stiffness was obtained with the addition of the hollow glass beads.

Here, the stiffness could be increased from 2162 MPa to 7550 MPa. The smallest increase in stiffness was obtained with the addition of the hollow glass beads. It can be assumed that the stiffness of the hollow spheres themselves is significantly lower than that of solid spheres, which means that the increase in Young's modulus is correspondingly lower.At the same time, however, the elongation at break of PA613 blends decreases sharply with increasing filler content. Even at a filler volume fraction of 20%, the elongation at break for blends with hollow glass beads (iM16K) and Cerabeads is below 4% in the dry state. The addition of fillers has a less pronounced effect on the strength of PA613 blends in the dry state.

The addition of fillers has a less pronounced effect on the strength of PA613 blends in the dry state. While the strength remains almost constant with the addition of glass beads, it decreases slightly with increasing filler content of hollow glass beads and Cerabeads.Comparing these results to the result after conditioning the samples, a massive impact of the moisture absorption especially for blends with PA613 and glass can be noticed. During conditioning, unfilled PA613 absorbs approx. 1.05 mass percent water. Depending on the filler, the blends absorb between 0.35 and 0.95 mass percent water. Due to the plasticizing effect of the water in PA613, the stiffness of the matrix decreases from 2162 MPa to 1507 MPa. The addition of 20 vol% Silibeads increases the modulus marginally to 1594 MPa. Increasing the filler content to 40 vol%, however, lowers the stiffness again. The addition of iM16K also causes only a slight increase in stiffness.

Increasing the filler content to 40 vol%, however, lowers the stiffness again. The addition of iM16K also causes only a slight increase in stiffness. The effect of conditioning on Cerabeads filled PA613 appears to be considerably less. Although the stiffness of these blends also decreases after conditioning, an increase in stiffness to 4850 MPa can be achieved with a filler content of 40 vol% Cerabeads.232Looking at the elongation at break after conditioning, it is first noticeable that there is an increase to almost 30% for unfilled PA613. This effect is also known and can be explained by the plasticizing effect of the water. The influence on the elongation at break of the blends is, however, even more pronounced. The elongation at break of PA613 with 40 vol% Silibeads with Si3 coating increases almost by a factor of 10 to 14% compared to the dry state.

The elongation at break of PA613 with 40 vol% Silibeads with Si3 coating increases almost by a factor of 10 to 14% compared to the dry state. The elongation at break of the blend with Silibeads without coating increases even further.Figure 11: Mechanical properties of PA613 based dry blends build in X-directionWhile the addition of 20 vol% Silibeads with Si3 coating had a positive effect on the strength of PA613 components, the strength of this blend drops significantly after conditioning. In233fact, the addition of 20 vol% Silibeads reduced the tensile strength of conditioned PA613 from 49.8 MPa to 31.2 MPa. As a simplification, it can be assumed that a filler content of 20 vol% reduces the cross-sectional area of the matrix by an average of 20 %. Due to the spherical shape of the fillers, however, this value fluctuates locally.

Due to the spherical shape of the fillers, however, this value fluctuates locally. Since the reduction in tensile strength due to the addition of 20 vol% glass beads is 35%, it can be assumed that the glass beads transmit almost no forces and that almost exclusively the matrix transmits the forces. The addition of 40 vol% glass beads deteriorates the tensile strength by approx. 60 %. There is no significant difference between uncoated glass beads and glass beads with silane coating. The tested silane coating therefore does not appear to be suitable for improving the adhesion between the glass beads and PA613 in laser- sintered components, especially since also no significant differences were found between the coated and uncoated glass beads. The tensile strength for blends with Cerabeads and iM16K decreased also after conditioning, but not to the same degree as the Silibeads. The adhesion was probably not degraded as much by conditioning.

The adhesion was probably not degraded as much by conditioning. Potentially, the surface structure of the Cerabeads help to maintain better adhesion by providing some mechanical interlocking.The only filler that can be classified as reasonable in combination with PA613 after the investigation of the mechanical properties are Cerabeads. At least if the components are to be used in a humid environment, dry blends of PA613 and glass lose their stiffness and strength. In addition to the plasticizing effect on the matrix, a degradation of the adhesion between polymer and filler was identified as the cause by analyzing SEM Images of the crack surfaces.The results for the PP based blends are displayed in figure 12. Although the DIN EN ISO 1110 standard is actually only intended for polyamides and PP hardly absorbs any water, the PP tensile specimens were also tested in the dry state and after storage in a climatic chamber at 70°C and 62% relative humidity for 7 days.

Even if the base polymer of the blend is hardly affected by the conditioning, the properties of the composite can still change if the fiber matrix adhesion is changed. Due to the poor processability of PP with Cerabeads as filler on EOS P3 systems, hardly any reliable data are available for this material combination.Unfilled PP is significantly more ductile compared to PA613. The strength in the dry state is 28.9 MPa and the modulus of elasticity is 1316 MPa. The material achieves a nominal elongation at break of 41%. Again, the stiffness in the dry state can be increased by adding any of the fillers tested and, as for PA613, increases with increasing filler content. The largest increase in stiffness was obtained with 40 vol% Silibeads with Si3 coating. However, the elongation at break decreases strongly with increasing filler content.

The largest increase in stiffness was obtained with 40 vol% Silibeads with Si3 coating. However, the elongation at break decreases strongly with increasing filler content. PP blends with a filler content of 20 vol% achieved good elongation at break between 8.7 % and 13.5 %, but they do not achieve the desired stiffness.In contrast to PA613 blends, the strength of PP blends decreases significantly with increasing filler content even in the dry state. With the addition of 20 vol% filler, the hollow glass spheres achieve the best strength. For both glass hollow spheres and glass solid spheres, the strength drops to approx. 17 MPa at a filler volume fraction of 40%. This corresponds to a decrease of 40 % in strength and indicates that only low forces can be transmitted by the filler. The filler reduces the effective force-transmitting cross-section of the matrix under tensile load accordingly. The strength of all PP blends is below the values of PA613 blends.

The filler reduces the effective force-transmitting cross-section of the matrix under tensile load accordingly. The strength of all PP blends is below the values of PA613 blends. It has been shown that the lower strength of the PP cannot be increased by the addition of spherical fillers. This would require fillers with a large aspect ratio, which can achieve a corresponding strengthening effect.234The results after conditioning show that PP blends are also affected by moisture even though hardly any moisture absorption could be measured (< 0.02%). The influence on component stiffness is not as significant as for PA613. The Young's modulus of the blend with 20 vol% Silibeads drops from 1910 MPa to 1620 MPa after conditioning.

The influence on component stiffness is not as significant as for PA613. The Young's modulus of the blend with 20 vol% Silibeads drops from 1910 MPa to 1620 MPa after conditioning. At the same time, the strength decreases from 20.7 MPa to 16.5 MPa and the elongation at break increases from 11.5% to 25%.Figure 12: Mechanical properties of PP based dry blends build in X-directionSEM images of the fracture surfaces of the tensile specimens are used to evaluate the adhesion between the filler and matrix. Ideally, the composite should fail in the matrix and not at the interfaces of filler and matrix. Accordingly, hardly any exposed filler without remaining plastic adhesions should be visible in the fracture surface. Exemplary, the fracture surfaces of PA613 with 20 vol% Silibeads with Si3 coating are investigated before and after conditioning (Figure 13).235Figure 13: SEM crack surface.

Exemplary, the fracture surfaces of PA613 with 20 vol% Silibeads with Si3 coating are investigated before and after conditioning (Figure 13).235Figure 13: SEM crack surface. PA613 + 20 vol% Silibeads Si3 coated tested dry (left) and after conditioning (right)The fracture surface after the tensile test in dry condition is shown on the left side. Only a few exposed glass spheres can be seen, and most glass spheres still have adherent plastic after the tensile test. The fracture surfaces of the polymer matrix show little deformation and an overall brittle fracture pattern. The glass beads are well embedded in the matrix, wetting with plastic in the LS process is consequently sufficiently good.After the specimens have been conditioned at 70°C and 62% relative humidity for 7 days, their fracture pattern changes fundamentally. There are more exposed glass spheres to which no more plastic adheres. A gap to the plastic is clearly visible at the edge of the glass spheres.

There are more exposed glass spheres to which no more plastic adheres. A gap to the plastic is clearly visible at the edge of the glass spheres. Since it could be concluded from the fracture surfaces of the specimens before conditioning that good wetting takes place in the LS process, it can be concluded that the plastic detaches from the glass during the tensile test. The glass spheres actually act as a defect. Obviously, almost no forces are transmitted by them. This also explains why the strength drops so significantly when glass beads are added in the conditioned state compared to the unfilled material. The polymer is stretched around the glass beads and has a more ductile failure mode. The silane coating of the glass beads does not seem to be suitable to ensure sufficient adhesion between matrix polymer and filler on laser sintered samples after conditioning until breakage.

The silane coating of the glass beads does not seem to be suitable to ensure sufficient adhesion between matrix polymer and filler on laser sintered samples after conditioning until breakage. One assumption for explaining this effect is that water diffusing into the polyamide collects at the interfaces between filler and polymer. Due to its polarity, it can detach dipole interactions between polyamide molecules and the glass surface.SummaryThe investigations showed that dry blending of spherical micro fillers into LS powder can influence the relevant processing properties. In addition to an increased melt viscosity, the change in the thermal properties of the powder is particularly relevant to the process. For this purpose, a concept was presented to adapt the laser power and the build chamber temperature specifically for different powder mixtures. Furthermore, it was shown that the component stiffness can be deliberately increased by the addition of spherical fillers.

Furthermore, it was shown that the component stiffness can be deliberately increased by the addition of spherical fillers. However, the elongation at break236decreases in the process. When testing conditioned tensile specimens, it was found that the mechanical properties of PA613 blends with glass in particular are strongly influenced by moisture absorption. Besides the plasticizing effect of moisture on the PA, a deterioration of the filler-matrix adhesion was identified as the cause. Adhesion between PA and Cerabeads was affected less by the influence of water. Accordingly, the silane coatings of the glass fillers could not effectively prevent the loss of adhesion. Further investigations on the improvement of the adhesion between filler and plastic in the LS process will be carried out at the DMRC in the future.References[1] Wohlers Associates, Ed., 3D Printing and Additive Manufacturing State of the Industry:Anual Worldwide Progress Report.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffects of Stripe Width on the Porosity and Mechanical Performance of AdditivelyManufactured Ti-6Al-4V PartsArash Soltani-Tehrani1,2, Mohammad Salman Yasin1,2, Shuai Shao1,2, Nima Shamsaei1,2,* 1Department of Mechanical Engineering, Auburn University, Auburn, AL, USA 2National Center for Additive Manufacturing Excellence (NCAME),Auburn University, Auburn, AL, USACorresponding author: Email: shamsaei@auburn.edu Phone: (334) 844 4839AbstractIn laser powder bed fusion (L-PBF) additive manufacturing, parts are manufactured in a layer-by-layer pattern. In each layer, cross-sections can be scanned with or without partitioning by the laser, which are commonly known as “stripe” and “meander” patterns, respectively.

In each layer, cross-sections can be scanned with or without partitioning by the laser, which are commonly known as “stripe” and “meander” patterns, respectively. Under the two scanning strategies, the thermal history experienced by the part can be considerably different. Accordingly, defect distribution, microstructure, and mechanical properties may be affected. In this study, two sets of Ti-6Al-4V specimens were fabricated using L-PBF: in one set, the tracks were partitioned in 5-mm stripes, while in the other set, they were partitioned at 100 mm resulting in no stripe seams in the cross-section. It was found that altering the stripe width can considerably affect the laser penetration depth, the defect content, and consequently fatigue performance.

It was found that altering the stripe width can considerably affect the laser penetration depth, the defect content, and consequently fatigue performance. However, tensile strength was not much sensitive to changing the stripe width.Keywords: Additive Manufacturing, Ti-6Al-4V, Stripe Width, L-PBF, Process ParametersIntroductionIn laser powder bed fusion (L-PBF) additive manufacturing, the part performance is dependent on the powder feedstock as well as design and process parameters.

For instance, it has been reported that if the powder is reused or its characteristics such as the size and shape are altered, the initial process parameter set may be insufficient to achieve a full melting, and consequently, result in the formation of a larger population of gas-entrapped pores or lack of fusion defects [1–5].With identical process parameters, altering the geometry alone can result in the parts experiencing different thermal histories [6–8] and as a result different defect content and mechanical properties. For example, it has been stated that the fatigue resistance of parts machined from additively manufactured (AMed) 17-4 PH SS large blocks were superior to the ones obtained from oversized dog-bone parts [9]. Moreover, complex geometries such as lattice structures have been reported to give different mechanical properties when the size of the features is altered [10– 12].

Moreover, complex geometries such as lattice structures have been reported to give different mechanical properties when the size of the features is altered [10– 12]. For instance, when the size of struts in lattices was decreased, a higher surface roughness, less ductility, and ultimate tensile strength were noted in Ref. [13].

For instance, when the size of struts in lattices was decreased, a higher surface roughness, less ductility, and ultimate tensile strength were noted in Ref. [13]. Therefore, adjusting the process parameters to achieve a consistent thermal history between different components, is essential to attain comparable part performance especially when complex geometries with different features are fabricated [10,14].1061Although many AM process parameters can influence the microstructure and defect content of parts, including but not limited to laser power, scanning speed, hatching distance, stripe width, layer thickness, laser spot size, inert gas flow speed and pressure, inert gas type, build plate preheating temperature [14–25], there only exist studies in the literature to quantify a few of them (such as laser power and scan speed). The precise effect of other parameters remains elusive.Stripe width is among the parameters whose effect on part performance is less well known.

The precise effect of other parameters remains elusive.Stripe width is among the parameters whose effect on part performance is less well known. As seen in Fig. 1(a), the stripes can partition the scanned area so that the laser tracks are not appreciably large. The shorter molten tracks are believed to result in smaller temperature gradients within each scan track (see Fig. 1(b)). Specifically, when the laser is moving along long tracks, the temperature at the onset of the track will be much less than the end of the track, resulting in a higher temperature gradient.(a)(b) Fig. 1 (a) Schematic illustration of the stripe width, hatching distance, and laser tracksduring the L-PBF process, and (b) temperature gradient in the short and long laser tracks.Therefore, this study aims to investigate the effects of fabricating specimens with and without striping for L-PBF Ti-6Al-4V (Ti64).

Following the introduction, the experimental program employed to characterize the part performance will be explained. Results will be subsequently presented in the order of defect content, microstructure, melt pool penetration depth, hardness, and tensile and fatigue properties. In addition, the observations on the experimental results will be discussed further, and lastly, some major conclusions will be drawn based on the results and discussions.1062Experimental ProgramIn this study, plasma-atomized Ti-6Al-4V (Ti64) from AP&C, a GE Additive Company with an initial particle size distribution (PSD) of 15-53 µm was used. Two sets of specimens were fabricated in one print in an argon environment using EOS M290, an L-PBF AM machine. Ti64 performance process parameters were used to fabricate all parts including the infill laser power (P) of 280 W, scanning speed (V) of 1200 mm/s, hatching distance (h) of 0.14 mm, and 30 µm layer thickness (t) resulting in 55.6 J/mm3 energy density.

All the process conditions for both sets of specimens were kept the same and the only difference was that one set was fabricated using a stripe width of 5 mm (hereon referred to as default) and the other fabricated with 100-mm stripe width (hereon referred to as modified) to ensure there would be no partition on the cross-section of the parts (see Fig. 2 for the layout). It needs to be specified that these two strategies are also known as “stripe” and “meander” scan patterns, respectively.(a)(b) Fig. 2 (a) Schematic illustration of the build layout that was used for L-PBF Ti64 fabrication, and (b) as-fabricated parts. Default and modified specimens are respectively enclosed with blue and red dashed boxes.On the build plate, there were 10 cylindrical specimens in each set. In addition, one specimen per condition was manufactured with a height half of the full cylinders for1063microstructural analyses to represent the gage section of the fatigue specimens.

In addition, one specimen per condition was manufactured with a height half of the full cylinders for1063microstructural analyses to represent the gage section of the fatigue specimens. These parts are illustrated as “half-built specimens” in Fig. 2(a). All specimens were located near the feedstock and center to minimize the effects of powder segregation on the mechanical properties [26]. All parts were manufactured on 3 mm support structures for straightforward removal of the build plate. After fabrication, the half-built specimens were removed in their non-heat-treated (NHT) condition, and the remainder were stress relieved (SR) on the build plate. The SR treatment was conducted at 900 °C for 1 hour in an argon environment followed by furnace cooling.All cylinders were machined to the geometry of fatigue specimens according to ASTM E466 (see Fig. 3) with a gage diameter of 0.2 inches (~5 mm) after SR [27].

3) with a gage diameter of 0.2 inches (~5 mm) after SR [27]. After machining, tensile tests were conducted on two specimens from each set on the same geometry with a procedure based on ASTM E8 [28]. Tensile tests were carried out in displacement-control mode with a nominal strain rate of 0.001 mm/mm/s in an MTS Landmark servohydraulic load frame equipped with a 100 kN load cell. Before tests, the gage length of specimens was marked to measure the percent elongation to failure (%EL) after the specimen fractures. In addition, only the first 0.05 mm/mm (5%) strain of each test was performed with an extensometer for an accurate measure of yield strength (YS). Force-controlled fully-reversed (R= -1) fatigue tests were also conducted using an MTS Bionix Tabletop load frame with a 25 kN load cell at two different stress levels according to ASTM E466 [27].Fig. 3 The geometry of fatigue specimens according to ASTM E466.

3 The geometry of fatigue specimens according to ASTM E466. All dimensions arein millimeters.Two specimens were randomly selected for X-ray Computed Tomography (XCT) analyses. The XCT was conducted via the Zeiss Xradia 620 Versa with a voltage, current, and voxel size of 140 kV, 150 µA, and 6 µm, respectively. After scanning, image slices were reconstructed using the built-in software. To ensure the reliability of the defect analysis and exclude the false defect detection from noise in the data, any defects smaller than 10 µm were discarded. The reason to exclude such defects was that small defects have usually less probability of fatigue cracks nucleation [14]. In addition, 3D visualizations of the defect distribution within the specimens were generated using DragonFly software [29].As previously mentioned, the half-built specimens were used for microstructural analyses.

At first, the specimens were cut in transverse and normal directions (TD and ND) and etched using the modified Krolls’ reagent (10% HF, 10% HNO3, and 80% distilled water). Afterward, the microstructure, melt pool boundaries and laser penetration depth were observed using a Keyence VHX-6000 digital microscope. In addition, hardness tests were conducted on the cross-section of half-build specimens via a LECO LCR 500 with a load rate of 60 µN/s. The hardness test results are presented in the Rockwell C (HRC) scale as suggested in ASTM E18 [30].1064Results and DiscussionsAs seen in Fig. 4(a) and (b), the microstructure in both specimens consists of an α+β Widmanstätten structure, which seems not to be affected by the stripe width. The α+β microstructure in the L-PBF Ti64 parts is mainly derived from the decomposition of α’ to α+β [31,32]. In addition, the prior β grains were found to be elongated toward the build plate as the heat is mainly dissipated in this direction.

In addition, the prior β grains were found to be elongated toward the build plate as the heat is mainly dissipated in this direction. The similar microstructure observed in each specimen can be because both specimens were manufactured using the same powder with comparable layer time intervals. Therefore, not much difference in the grain structure is expected.As labeled on the image, the melt pool depth was also calculated by measuring a minimum of 10 melt pools per condition, which showed 210 ± 24 µm and 156 ± 5 µm laser penetration depth for the default and modified conditions, respectively. The changes in laser penetration depth can be correlated with stripe width which should have resulted to some extent different thermal histories. It seems that when shorter tracks were used (i.e., default), there was excessive heat input especially at laser turning spots [33].

It seems that when shorter tracks were used (i.e., default), there was excessive heat input especially at laser turning spots [33]. The higher heat input should have resulted in deeper melt pools, which were on average ~35% larger for the default condition as compared with the modified one.(a)(b) Fig. 4 The microstructure and melt pool shapes in L-PBF Ti64 fabricated via (a) default and (b) modified stripe width in their NHT conditions for both transverse and normal directions (TD and ND). The arrows show the melt pool depth along the building direction.As seen in Fig. 5, the defect distributions in both specimens are shown with 3D visualizations. It seems that there are a larger number of defects when the default scan pattern was used. However, some large defects are also evident in the case of the modified scan pattern. Observing more defects in the default specimen was quite interesting.

However, some large defects are also evident in the case of the modified scan pattern. Observing more defects in the default specimen was quite interesting. As previously explained,1065smaller stripes are supposed to result in a uniform temperature distribution along the tracks. However, it seems that the presence of more turns and excessive energy input at the beginning and end of shorter tracks (as the laser slows down to turn) can result in the formation of keyhole pores in such locations. This observation was consistent with Ref. [33] where a higher probability of keyholes was reported in laser turning spots in serpentine strategies.The observations of more defects near the stripe boundaries and in some cases keyholes have been attributed to the fast cooling rates in the L-PBF process, having less time for the molten material to fill in the empty spaces [34].

The authors are aware of an option known as “skywriting” in the EOS M290 which supposedly controls the laser turning to occur outside of the hatch lines to prevent overheating at the onset and end of each track. However, Martin et al. [33] stated that skywriting still cannot alleviate this issue due to the physical limitations of mirror-based scanning setups. Alternatively, varying laser power along the tracks was introduced as the efficient solution to make the energy density uniform at turn spots [33].Tensile stress-strain and stress-displacement curves are reported in Fig. 6. As seen in this figure, tensile properties including the ultimate tensile strength (UTS), yield strength (YS), and percent elongation to failure (%EL) were almost comparable for both cases. The UTS was ~1013- 1014 MPa for both conditions. The YS was somewhat higher in the modified specimens, which can be attributed to slightly finer lath martensite thickness due to higher cooling rates in these specimens.

The YS was somewhat higher in the modified specimens, which can be attributed to slightly finer lath martensite thickness due to higher cooling rates in these specimens. Additionally, the %EL was almost comparable for both scan patterns. In general, consistent with the similar microstructures in both specimen sets, similar tensile properties were observed.Volume (µm3)DefaultModified177,000133,00068,75044,600525Fig. 5 3D visualizations of defect content in L-PBF Ti64 specimens manufactured usingdefault and modified scan patterns.1066(a)(b) Fig. 6 (a) Engineering stress-engineering strain and (b) engineering stress-displacement curves for L-PBF Ti64 parts fabricated using default and modified patterns in machined surface conditions.Hardness tests were also conducted for both default and modified specimens. As seen in Fig. 7, it was found that there is no difference in the HRC results.

As seen in Fig. 7, it was found that there is no difference in the HRC results. The similar HRC well correlates1067with the similar tensile properties and microstructure that were previously discussed. Therefore, it may be overall assumed that changing the stripe width does not alter the microstructure.Fig. 7 Rockwell C hardness (HRC) values for both default and modified L-PBF Ti64specimens in their NHT conditions.Fully-reversed fatigue tests were conducted on the specimens fabricated using the default and modified scan patterns and reported in Fig. 8. Tests were conducted at two stress levels (i.e., 700 and 600 MPa) to include the mid-cycle and high-cycle fatigue regimes. It was noticed that using a modified scan pattern can result in somewhat higher fatigue performance at lower stress levels (i.e., high-cycle fatigue). However, as the stress level was increased, the difference between fatigue lives was diminished.

However, as the stress level was increased, the difference between fatigue lives was diminished. This behavior can be explained by the fact that at lower stress levels, the majority of fatigue life is spent at crack nucleation whereas, at higher stress levels, it is mainly spent at crack growth [35].As the crack growth behavior is more dependent on the microstructure and no difference was noted in the microstructure of the parts fabricated using default and modified scan patterns (see Fig. 4), negligible variation between fatigue lives can be expected. The high-cycle fatigue regime, however, is typically governed by the defect content in the AMed parts. As a result, it may be assumed that critical defects resulting in fatigue failures were typically smaller in the specimens manufactured using a modified scan pattern.

As a result, it may be assumed that critical defects resulting in fatigue failures were typically smaller in the specimens manufactured using a modified scan pattern. As previously explained, due to excessive energy inputs at the onset and end of laser tracks in default specimens, some large pores can be present which are detrimental to fatigue resistance. A summary of all fatigue tests is also reported in Table 1.1068Fig.

A summary of all fatigue tests is also reported in Table 1.1068Fig. 8 Fatigue performance of L-PBF Ti64 parts fabricated using default and modifiedscan patterns in machined surface condition.Table 1 A summary of fully-reversed force-controlled fatigue tests for L-PBF Ti64 specimens in machined surface condition and manufactured using default and modified scan patterns.Specimen ID 05D 03D 09D 02D 04D 07D 08D 18M 20M 14M 21M 19M 16M 15M𝜎𝜎𝑎𝑎(MPa) 600 600 600 600 600 700 700 600 600 600 600 600 700 7002Nf (Reversals) 1,513,966 703,184 121,602 284,200 265,520 165,472 73,312 1,454,378 4,687,934 4,615,468 1,458,260 363,442 111,772 91,796Scan PatternDefaultModifiedThe fracture surfaces of some specimens were investigated to compare the defect sizes resulting in fatigue failures. It needs to be specified that all fracture surfaces were investigated however only two of them are presented here. As seen in Fig. 9, one specimen from each set is shown.

It needs to be specified that all fracture surfaces were investigated however only two of them are presented here. As seen in Fig. 9, one specimen from each set is shown. It can be seen that the specimen fabricated using the modified scan pattern had higher1069fatigue life. The equivalent defect sizes were evaluated based on Murakami’s approach which considers the √area of the defect [36,37]. It can be noted that the default specimen had a smaller crack initiation site however it was at the surface resulting in considerably smaller fatigue life. In the case of modified, the defect was away from the surface.(a)(b)Fig. 9 Fracture surfaces of L-PBF Ti64 specimens manufactured using (a) default and (b)modified scan patterns. Both specimens were tested at 600 MPa.To have a better estimate of all the critical defect sizes that resulted in fatigue failures, the √area of critical defects was plotted on a Gumbel curve. As reported in Ref.

As reported in Ref. [38], the critical defect size obtained by √area follows the extreme value distribution which is shown in Fig. 10. The reduced variate (Y), shown in this figure, is calculated by Y= -ln(-ln(i/N+1)), where i is the rank of the defect based on its size and N is the total number of defects investigated [39]. From this1070figure, it can be noticed that for a majority of tests, there is a significant overlap in defect sizes between the two sets. However, for some specimens, some larger defects were also observed in the case of default scan pattern whereas smaller defect sizes were seen in the counterpart specimens manufactured using modified scan pattern.Fig.

10 Extreme value distribution for L-PBF Ti64 specimens manufactured usingdefault and modified scan patterns.ConclusionsIn this study, two sets of L-PBF Ti64 specimens were fabricated by utilizing different stripe widths during scans, one with a 5-mm stripe as default, and another one with 100 mm as the modified (also known as meander scan pattern). Following conclusions can be drawn based on the experimental results:Using a smaller partition size may result in larger laser penetration depth due to excessive energy input in laser turn spots.The number of defects increased with decreasing the stripe width which was attributed to more frequent laser turn spots.Stripe width variations resulted in negligible changes in the microstructure, HRC, and tensile properties.Fatigue performance was to some extent improved at lower stress levels (i.e., high- cycle fatigue regime) when stripe width was enlarged.

However, as the stress level was increased, the difference between fatigue lives was diminished.In summary, it seems that altering the stripe width can considerably affect the defect content of AMed parts. This difference can be even further pronounced when more complex geometries are fabricated. Therefore, more investigation into the effects of stripe width is indeed1071recommended. It is also important to study the effects of stripe width in large components where residual stress during fabrication might be a challenge.AcknowledgmentThis material is based upon work partially supported by the National Institute of Standards and Technology (NIST) under Award No. 70NANB19H170 and National Science Foundation (NSF) under Grant No. 1919818.References[1]P.E. Carrion, A. Soltani-Tehrani, S.M. Thompson, N. Shamsaei, Effect of Powder Degradation on the Fatigue Behavior of Additively Manufactured As-Built Ti-6Al-4V, Solid Free. Fabr. Proc. (2018).[2] A. Soltani-Tehrani, J.

FECTS OF TEMPERATURE ON AQUEOUS FREEFORM EXTRUSION FABRICATIONJie Li1, Ming C. Leu1 and Gregory E. Hilmas2 1Department of Mechanical and Aerospace Engineering 2Department of Ceramic Engineering Missouri University of Science and Technology, Rolla, MO 65409AbstractAn experimental study was conducted to investigate the effect of temperature on ceramic parts produced by paste extrusion based additive manufacturing followed by sintering. A computer-controlled gantry system equipped with a piston extruder was used to extrude aqueous alumina paste. The system includes a temperature control subsystem that allows for freeform extrusion fabrication inside a low-temperature (<0°C) chamber. It can also be used for fabricating parts on a hot plate at ambient or higher temperatures (≥20°C). Test specimens were fabricated from aqueous aluminum pastes at -20°C in the low-temperature chamber and also on the hot plate at 40°C.

Test specimens were fabricated from aqueous aluminum pastes at -20°C in the low-temperature chamber and also on the hot plate at 40°C. The minimum angles achievable by these two processes for part fabrication, without use of support material, were compared. Also compared were the relative density and mechanical properties of the parts obtained after sintering. Microstructures were examined via scanning electron microscopy in order to obtain a deeper understanding of the effect of fabrication temperature.IntroductionSince the mid-1980s, Solid Freeform Fabrication (SFF) technology has been developed with the potential of becoming an efficient and inexpensive manufacturing technique to produce polymer, metal and ceramic parts in small volumes [1, 2].

SFF techniques for ceramic component fabrication include Ink-jet Printing [3], Stereolighography (SLA) [4], 3D Printing (3DP) [5, 6], Selective Laser Sintering (SLS) [7], Robocasting [8] and Fused Deposition of Ceramics (FDC) [9, 10]; see Table 1. Most of SFF techniques for ceramic component fabrication involve the use of high concentration (>40%) of organic binders that needs to be removed during post processing, which generate harmful wastes for the environment. One of extrusion deposition techniques of ceramic is FDC, which is capable of fabricating dense ceramic parts with good surface accuracy but uses a relatively large amount (40-50%) of organic chemicals as binders [10].Use of an aqueous process with a much lower binder amount would be more environmentally friendly. Robocasting is one well-known technology, initially developed at Sandia National Laboratories, for fabrication of ceramics and ceramic composites.

Robocasting is one well-known technology, initially developed at Sandia National Laboratories, for fabrication of ceramics and ceramic composites. This process could extrude 50~65% high solids loading aqueous slurry with less than 1% organic binder. For fabricating solid, dense parts, Robocasting uses a heating source (a 40oC hot plate) to dry the paste during extrusion to form a 3D part [8]. The relative density and flexural strength reported were 93.7% and 310 MPa for Al2O3 samples [8, 11].Another aqueous SFF technology is Freeze-form Extrusion Fabrication (FEF), which was developed by researchers at Missouri University of Science and Technology. In the FEF process, a high solids loading (>50%) aqueous paste containing 1~4 vol% organic additives is deposited inside a freezing chamber (<0oC) to solidify the paste after it is extruded. Freeze drying is thenused to prevent crack formation during the process of removing water content.

Freeze drying is thenused to prevent crack formation during the process of removing water content. The flexural strength reported was 219 MPa for Al2O3 samples [12, 13].Table 1 Popular SFF techniques for ceramic material Process Selective Laser Sintering (SLS)Method Sinter ceramic-binder mixed in powder bedMaterials Al2O3, SiC, ZrSiO4References [7]3D printing (3DP)Print binder solution on ceramic powder bedAl2O3, Si3N4[5, 6]Stereolighography (SLA)Cure mixed resins with ceramic particlesSiO2, PZT, Al2O3[4]Ink-jet PrintingPrint colloidal dropletsAl2O3, Si3N4, ZrO2[3]Fused Deposition Ceramic (FDC)Print melt ceramic particle-filled polymerAl2O3, Si3N4,[9, 10]Robocasting (In air/oil)Print organic/non-organic concentrated colloidal gelAl2O3, PZT, SiO2[8, 11]This research systematically studied the effect of temperature on the mechanical properties of freeform extrusion fabricated ceramics from aqueous pastes.

A custom-built 3D gantry system was equipped with a cooling sub-system and could include a hot plate for part fabrication. This system was used to print aqueous alumina paste by freeform extrusion fabrication on a 40°C plate inside a room-temperature chamber and on a -20°C plate inside a -20°C temperature chamber, in order to study the effect of temperature on fabrication results. In this paper, the first condition (fabricating on a 40°C plate in a room-temperature environment) would be referred to ‘at 40oC’, and the second condition (fabricating on a -20°C plate inside a -20°C temperature chamber) would be referred to ‘at -20oC’. A 60% solids loading aqueous Al2O3 paste was used as the part material in all of the experiments. The relative density, mechanical properties, part accuracy, and minimum deposition angle of Al2O3 parts fabricated at different temperatures were investigated.

The relative density, mechanical properties, part accuracy, and minimum deposition angle of Al2O3 parts fabricated at different temperatures were investigated. SEM and optical images were used to help understand the temperature effect on the microstructure of the fabricated parts.Experimental Setup and MethodsMachine OverviewThe experimental system consists of a motion subsystem, a control subsystem, and extrusion devices. A picture of the overall system is shown in Fig. 1a. The system has three linear axes (Parker Hannifin, Daedal 404 XR) driven by three stepper motors (Empire Magnetics). In this research, a single extruder is used to extrude aqueous alumina paste. The paste is extruded onto a substrate that moves in the x and y axes. After deposition, the paste would solidify in a freezing environment as shown in Fig. 1b or dry on a hot plate set at 40oC as shown in Fig. 1c.When the fabrication of one layer is completed, the gantry moves up by one layer thickness.

1b or dry on a hot plate set at 40oC as shown in Fig. 1c.When the fabrication of one layer is completed, the gantry moves up by one layer thickness. These processing steps are repeated until the entire part is formed.(a)(b) (c)Figure 1. Experimental setup of the machine: (a) overview of machine, (b) the cooling system and (c) the hot plateMaterial CharacterizationThe 60% solids loading alumina paste was used in all of the experiments in this study. The particle size were analyzed using a Microtrac Particle Size Analyzer (S3500, Microtrac), and the particle surface area was measured using a NOVA 2000e instrument (Quantachrome Instruments). The paste consisted of a combination of Al2O3 powder (A-16SG, ALMATIS, 0.34 µm, 9.44 m2/g), glycerol (Aldrich), DARVAN® C-N (ammonium polymethacrylate, Vanderbilt Minerals), Methocel*F4M (methylcellulose, Dow Chemical) and deionized water.

The slurry was mixed with Darvan C and glycerol and subsequently ball-milled for 10 hours to break up agglomerates and produce a uniform mixture. Darvan C with a negative surface charge was used as a dispersant to mitigate the Van Der Waals forces between particles [14-16]. Glycerol (20 wt%) was used to prevent the growth of large ice crystals and freezing defects associated with water solidification [17]. Methocel was dissolved in water at 70°C after 5 minutes of mechanical stirring to form a 60 vol% solids loading paste and was chosen as a binder to increase the paste viscosity and assist in the formation of a stronger green body after drying. A vacuum mixer (Whip Mix, Model F) was used to remove air bubbles by degassing for 10 minutes.Post-processingThe water content in the test bars fabricated at -20°C was removed via sublimation using a freeze dryer (Virits, Model Genesis 25XL).

When building the part on a hot plate, the water inside the paste evaporated at 40°C to increase solids loading, which provides the strength for layer-by-layer fabrication to form 3D parts.The results of thermogravimetric analysis (TGA) and sintering tests were used to determine the debinding and sintering schedule. After drying, the samples were pyrolyzed to remove the remaining organics using a 0.5°C/minute ramp up to 500°C with a hold of 2 hours. Next, the samples were sintered to 1,550°C using a heating rate of 10°C/minute, held for 2 hours, and then cooled to room temperature.Mechanical Properties and MicrostructureThe fabricated samples, after sintering, were ground by diamond grinding according to ASTM C-1161 standard “A” bars (2x1.5x20 mm3). Four-point bending tests were performed on a screw-driven mechanical frame (Instron, Model 5881) to measure the flexural strength and elastic modulus. A micro-hardness tester (Struers, Model Duramin-5) was used to measure hardness.

A micro-hardness tester (Struers, Model Duramin-5) was used to measure hardness. The hardness was measured using a load of 1 kg, and five measurements were collected for each specimen on a 0.25 μm diamond polished surface.Bars fabricated at 40°C, -20°C, and -20°C without using a nozzle for extrusion. Three bars were printed to test the green body density. After drying and binder removal, Archimedes method (in water) was used to measure the green body density and relative density after sintering.The microstructures of fabricated bars were studied and compared in order to understand the effect of printing flaws and formation of ice crystal voids. Samples were polished to a 0.25 μm surface finish, and SEM was used to examine the microstructure of each sample.Results and DiscussionOverall Effects of TemperatureWhen the paste printing was done on a plate (without heating) at room temperature without using a hot plate, the drying rate was not fast enough.

In this case, a large 3D part being fabricated will deform or even collapse. In contrast, the drying rate of the extruded material on a 60oC plate (in a room temperature environment) was significantly higher than printing on a 40oC plate. Moisture distributed unevenly in the part’s body as a result of the temperature variations and non- uniform drying led to part warping or cracking. Thus, 40oC was used for the hot plate temperature. Ten bars (6x7x60mm3) were fabricated at 40°C and ground to the standard B bar size. The average flexural strength of bars fabricated at 40°C was 253 MPa and the average elastic modulus was 327 GPa as listed in Table 2. These bars have lower flexural strength than pressed bars (370~390 MPa). A few printing flaws and air bubbles in the paste caused ~6% porosity, which markedly reduced mechanical strength after sintering. Ten additional single walls having only one filament at each layer were fabricated in an attempt to eliminate printing flaws.

Ten additional single walls having only one filament at each layer were fabricated in an attempt to eliminate printing flaws. The difference between the nozzle printed bar’s relative density and the flaw-free single wall’s relative density was smaller than 1% (see Table 2). Thus, the air bubble in the paste was the main reason for the approximately 3% porosity of parts fabricated at 40oC.A heating coil (see Fig. 1a) was used to maintain the paste warm when printing was done in the freezing chamber. The extruded ceramic paste could freeze at -20°C and the required waiting time for one filament (0.5*0.74*60 mm3) was approximately 10 seconds. One critical issue inprinting under a freezing environment was clogging, which often occurred because the paste froze inside the nozzle before it was extruded.Six bars (6x7x60 mm3) were fabricated at -20°C and ground to the standard B bar size.

The average flexural strength of bars fabricated at -20oC was 153 MPa, and the average elastic modulus was 327 GPa (see Table 3). Ice crystals, printing flaws, and air bubbles were primarily responsible for this kind of low flexural strength. Bars fabricated at -20°C tended to have more flaws than those fabricated at 40°C. Two factors were responsible for this increased number of flaws. First, clogging led to discontinued printing and voids. Second, the paste extruded at -20°C solidified faster than drying of the paste extruded at 40°C, leaving the time required to fill the voids between filaments too short. Single walls were also printed to measure the part density without printing flaws, and it indicated that ice crystal voids also contributed to the lower density.Table 2.

Single walls were also printed to measure the part density without printing flaws, and it indicated that ice crystal voids also contributed to the lower density.Table 2. Mechanical properties and relative density of bars fabricated at 40°C# Flexural Strength (MPa)1 2 3 4 5 6 7 8 9 10 Average Standard Deviation327 290 276 268 262 246 244 234 207 177 253 42Flexural Modulus (GPa)352 333 341 330 349 320 286 316 306 337 327 20Relative Density (%)94.40 96.30 92.20 92.62 94.03 94.14 92.24 94.04 93.67 94.15 93.78 1.22Relative Density of Single Walls (%) 93.63 95.01 94.81 94.73 94.47 94.86 94.55 94.34 94.32 96.85 94.76 0.83Table 3.

Mechanical properties and relative density of bars fabricated at -20°C Relative Density of Single Walls (%) 86.52 87.58 87.30 84.67 86.36 86.71 86.52 0.93# Flexural Strength (MPa)Flexural Modulus (GPa)Relative Density (%)1 2 3 4 5 6 Average Standard Deviation192 166 161 154 132 116 153 24242 217 274 244 246 197 237 2488.74 86.91 88.83 88.06 89.12 85.74 87.90 1.21Two bars were fabricated on a 0oC plate inside a -20°C chamber. The relative density and flexural strength of the bars were 86.5% and 48 MPa, respectively. The bars’ mechanical properties were much lower than printing on a -20oC plate because large-sized voids formed from ice crystals as illustrated in Fig. 2a.Five ‘big’ bars were printed at -20°C directly from a syringe without using a nozzle to avoid printing flaws. Each of these ‘big’ bars contained only one filament, and the size was approximately 10x10x60 mm3 after deposition. After post-processing, the sample was ground to the standard B bar size.

Each of these ‘big’ bars contained only one filament, and the size was approximately 10x10x60 mm3 after deposition. After post-processing, the sample was ground to the standard B bar size. The ‘big’ bars did not contain any printing flaws, but the flexural strength and the elastic modulus of these bars (listed in Table 4) were even lower than the bars printed with a nozzle at -20oC (Table 3). The low strength is thought to be due to the relatively large filament having an uneven temperature gradient inside the filament. The temperature inside the filament and filament’s surface temperature were measured using a thermometer. As shown in Fig. 3, the temperature inside the big filament was warmer than the filament’s surface temperature. Ice crystal formation was observed as shown in Fig. 2b. Apparently larger-sized crystals formed when the temperature is warmer (between 0oC and -20oC).Table 4.

Ice crystal formation was observed as shown in Fig. 2b. Apparently larger-sized crystals formed when the temperature is warmer (between 0oC and -20oC).Table 4. Mechanical property and relative density of bars fabricated at -20°C without using a nozzle# Flexural Strength (MPa) 122 104 77 69 60 86 231 2 3 4 5 Average Standard DeviationFlexural Modulus (GPa) 186 210 179 139 149 173 26Relative Density (%) 86.19 85.23 85.01 86.07 85.46 85.59 0.46Ice crystal voids(a) (b)Figure 2. Cross-section view (a) bar fabricated on a 0oC plate and (b) bar fabricated at -20°C without using a nozzleFigure 3. Non-uniform temperature inside the filament fabricated at -20°C without using a nozzleFigure 4 provides a comparison of flexural strengths and relative densities for bars fabricated under three different conditions: 40oC, -20oC, and -20oC without using a nozzle.

The relative density of bars fabricated at 40°C was 94.76%, which was higher than the bars fabricated at -20°C with/without using a nozzle (86.52% / 85.59% in relative density). The relative density of bars fabricated at -20°C with/without using a nozzle had only small difference (1%). However, the flexural strength of bars fabricated at -20°C without using a nozzle was 40% lower than the flexural strength of bars fabricated at -20°C with using a nozzle (see Fig. 4a).(a) (b)Figure 4. Comparison of (a) mechanical properties and (b) relative density of parts fabricated by three different methodsThe hardnesses of samples fabricated at 40oC and -20oC were 16.78±0.43 GPa(1712.04±44.06 Kg/mm2) and 14.36±0.85 GPa (1465.46±86.26 Kg/mm2), respectively.Part Accuracy and Minimum Deposition AngleBars were printed at 40°C and -20°C to compare part accuracy as shown in Fig. 5. The dimensions were measured for the green (after drying) and sintered bars, as listed in Table 5.

5. The dimensions were measured for the green (after drying) and sintered bars, as listed in Table 5. Bars fabricated at -20°C had 1.5% to 3% more expansion than bars fabricated at 40°C because each filament contained approximately 40 vol% water and freezing of water increased the total volume. The minimum deposition angle here refers to the angle that can be achieved between the substrate and the slope of a hollow cone without collapsing, as illustrated in Fig. 6a. This angle reflects the capability of a freeform extrusion fabrication process to build 3D parts without need of support material [18]. Two sets of tests were conducted to fabricate cones with different bottom diameters to find the minimum deposit angle. In each set of tests, hollow cones were fabricatedusing the bottom diameters of 38 mm, 51 mm and 64 mm.

In each set of tests, hollow cones were fabricatedusing the bottom diameters of 38 mm, 51 mm and 64 mm. The cone angle was varied from 60o to 20o by 5o decrements to measure the angle at which the cone collapsed and then the angle was increased from the failure angle by 2o increments to determine the minimum deposition angle. The part was printed with a 6 mm/s table speed. As shown in Table 6, the parts printed at -20oC had a smaller deposition angle than printed at 40oC. This is because each extruded filament would solidify faster at -20oC, providing support strength to prevent part from collapsing.(a) (b)Figure 5. Bars fabricated at (a) 40oC and (b) -20oCTable 5.

Bars fabricated at (a) 40oC and (b) -20oCTable 5. Shrinkage in the green body and sintered partCADGreen barW (mm) 5.6 6.03 ±0.12H (mm) 7.8 8.03±0.13L (mm) 70 71.12±0.12Printed at 40oCLinear shrinkage of green bar After sintering7.68±2.06% 5.05±0.16Linear shrinkage after sintering 9.88±2.90%2.91±01.64% 7.24±0.031.60±0.18% 60.54±0.16 13.51±0.23%1.60±0.18% 60.54±0.16 13.51±0.23%Green bar6.17±0.108.19±0.0272.37±0.20Printed at -20oCLinear shrinkage of green bar10.24±1.69%After sintering Linear shrinkage after sintering 10.06±2.70%5.04±0.155.04±0.26%7.00±0.11 10.21±1.39%3.39±0.29%58.78±0.21 16.02±0.29%(a) (b) (c)Figure 6. Hollow cone: (a) θ is the minimum deposition angle, D is the bottom diameter and H is the height of cone, (b) fabricated cone at 40oC and (c) fabricated cone at -20oC [18].Table 6.

Hollow cone: (a) θ is the minimum deposition angle, D is the bottom diameter and H is the height of cone, (b) fabricated cone at 40oC and (c) fabricated cone at -20oC [18].Table 6. Minimum deposition test resultsBottom Diameter = 38 (mm)Bottom Diameter = 51 (mm)Bottom Diameter = 64 (mm)Minimum deposition angle θ (°)Fabrication at 40oC505255Fabrication at -20oC [18]282624Mechanical PropertiesTen standard A bars were fabricated at both 40°C and -20°C and were subjected to the same post-processing. The bars fabricated at 40°C and -20°C had a green density of 58.49% and 51.35%, respectively. As shown in Tables 7 and 8, bars fabricated at -20°C achieved an average strength of 300 MPa, whereas the average strength of bars fabricated at 40°C was 338 MPa.The theoretical flaw size was calculated by using equation (1), which is Griffith criterion.

Assuming the fracture toughness to be 4 MPa*m1/2 [19], and a shape factor that is characteristic of joined particles, Griffith criterion is:𝜎𝑓 =𝐾𝐼𝐶 𝑌√𝑐(1)where 𝐾𝐼𝐶 is the fracture toughness, Y= π1/2 and c is one half of the maximum flaw size. Based on the results of standard A bar test (see Tables 7 and 8), the calculated and measured maximum flaw size for bars fabricated at -20oC and 40oC are given in Table 9. Example flaws for bars fabricated at -20oC and 40oC are illustrated in Fig. 7. As shown in Fig. 7a, the bars fabricated at -20oC have printing flaws (in triangular shape) in the cross section. These flaws may lead to lower strength in the printed bars than those fabricated at 40oC. On the other hand, bars fabricated at 40oC have smaller flaws than those fabricated at -20oC or even have no flaws such as samples #1 and # 2 in Table 8. But a few of big flaws may occur when fabricating is done at 40oC, such as sample #5 shown in Table 8, Table 9 and Fig.

But a few of big flaws may occur when fabricating is done at 40oC, such as sample #5 shown in Table 8, Table 9 and Fig. 7b, which leads to a larger standard deviation in the mechanical properties.Table 7. Mechanical properties and relative density of standard A bars fabricated at -20°C# Flexural Strength (MPa) 334 325 303 269 268 300 281 2 3 4 5 Average Standard DeviationFlexural Modulus (GPa) 261 259 356 285 240 280 40Relative Density (%) 90.48 93.32 91.34 91.41 91.18 91.55 0.95Table 8. Mechanical properties and relative density of standard A bars fabricated at 40°C# Flexural Strength (MPa) 406 361 348 346 232 338 571 2 3 4 5 Average Standard DeviationFlexural Modulus (GPa) 273 316 308 353 306 311 25Relative Density (%) 97.18 97.11 96.15 97.31 95.92 96.73 0.58Table 9.

Calculated and measured maximum flaw size for the standard A barStandard A BarBars Fabricated at 40°CBars Fabricated at -20°CCalculated Maximum Flaw Size (μm) Bar #3: 84 Bar #4: 85 Bar #5: 189 Bar #2: 96 Bar #3: 110 Bar #4: 141Measured Maximum Flaw Size (μm) Bar #3: 54 Bar #4: 84 Bar #5:177 Bar #2: 66 Bar #3: 91 Bar #4: 116Printing flawsSingle large flaw inside the bar(a) (b)Figure 7. Cross-section of (a) bar fabricated at -20°C and (b) bar fabricated at 40°CThe theoretical modulus could be calculated using Nielsen’s relationship of elasticity forporous ceramic materials as follows:E = E0(1−P)2 1 ρ1+(−1)P(2)where E0 is the pore-free elastic modulus, P is the porosity in volume percent and ρ is Nielsen’s shape factor (0.4).

Based on the standard A bar measured results in Tables 7 and 8 and assuming a modulus of 380 GPa [20] for fully dense (100% density) Al2O3, the bars fabricated at 40°C had 3.27% on average, and the bars fabricated at -20°C had 8.45% porosity. By using Eq. (2), the theoretical modulus calculated for bars fabricated at 40°C was 339 GPa, and for the bars fabricatedat -20°C was 282 GPa. The experimentally measured flexural moduli in Tables 7 and 8 agreed fairly well with the theoretically calculated values.Microstructure and Ice Crystal FormationThe SEM image of bars fabricated at -20°C revealed that several cracks spanned the entire bar (Fig. 8a). These cracks occurred because ice crystals formed at the boundary of each filament during freezing. Then the boundaries of neighboring filaments were not bonded strongly. Fig. 8b shows an image of the side of a bar fabricated at -20°C, and this image indicates ice crystal voids.

Then the boundaries of neighboring filaments were not bonded strongly. Fig. 8b shows an image of the side of a bar fabricated at -20°C, and this image indicates ice crystal voids. Bars fabricated at -20oC without using a nozzle had a warmer temperature (~ -15oC) inside the filament before they were fully frozen. The warmer temperature led to more formation of ice crystal voids [21, 22]. Figure 9 shows that the voids were larger and there were more voids than bars fabricated at -20oC with using a nozzle.The SEM of bars fabricated at 40oC shown in Fig. 10 showed that there might be air bubbles present inside the bar, and the overall microstructure was similar to that of pressed pellets.CrackPrinting flawIce crystal voids(a) (b)Figure 8. Bars fabricated at -20oC (at 350x zoom): (a) cross-section view and (b) side viewContinuous poreIndividual poreFigure 9.

Bars fabricated at -20oC (at 350x zoom): (a) cross-section view and (b) side viewContinuous poreIndividual poreFigure 9. Bars fabricated at -20oC without using a nozzle (at 350x zoom): (a) side view and (b) cross-section viewAir bubble void(a) (b) (c)Figure10. Cross-section of bars fabricated at 40oC (a) 100x zoom and (b) at 3000x zoom ;(c) cross-section of pressed pellet 3000x zoomConclusionThis paper discusses the effects of temperature on ceramic parts produced by freeform extrusion fabrication. The parts fabricated at 40°C achieved relative density, flexural modulus, and flexure strength of 96.73%, 311 GPa, and 338 MPa, respectively. The parts fabricated at -20°C attained the relative density, flexural modulus and flexure strength of 91.55%, 280 GPa and 300 MPa, respectively. At 40°C, the minimum deposition angle achieved was 50o at 38 mm bottom diameter, and at -20°C the minimum deposition angle achieved was 28o at the same bottom diameter.

At 40°C, the minimum deposition angle achieved was 50o at 38 mm bottom diameter, and at -20°C the minimum deposition angle achieved was 28o at the same bottom diameter. The test bars fabricated at 40oC and at -20oC both have high flexural strength, Young’s modulus, and relative density. The test bars fabricated at 40oC have slightly higher mechanical properties. The hollow cones fabricated at -20oC have smaller minimum deposition angles than those fabricated at 40oC, indicating the greater capability of a freezing environment for fabricating larger and more complex parts without use of support (sacrificial) material. The analysis of the SEM images of the fabricated parts helps understand the formation of ice crystals when fabricating parts by the freeform extrusion fabrication process at freezing temperatures.References[1] Jacobs, P. F., Rapid prototyping and manufacturing: fundamentals of stereolithography, 1st Ed., Society of Manufacturing Engineers (1992).

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffects of Thermal Camera Resolution on Feature Extraction in Selective Laser MeltingXin Wang, Cody S. Lough, Douglas A. Bristow, Robert G. Landers, and Edward C. Kinzel Department of Mechanical and Aerospace Engineering, Missouri University of Science andTechnology, Rolla, MO 65409AbstractSelective Laser Melting (SLM) is a common additive manufacturing process which uses a laser energy source to fuse metal powder layer by layer. Engineering properties and microstructure are related to the part’s thermal history. It is important to measure the thermal history in-situ to qualify parts and provide the sensing which is necessary for process control.

It is important to measure the thermal history in-situ to qualify parts and provide the sensing which is necessary for process control. A common measurement tool for this purpose is a thermal camera that records the thermal emission of the part’s surface.This study investigates the effects of spatial sampling resolution of thermal cameras when monitoring the temperature in SLM processes. High-fidelity simulation of an SLM process is used to quantify the effects of the camera’s sampling in space. Next, the effect that spatial resolutions have on feature extraction, namely peak temperature and melt pool morphology, is investigated by applying feature extraction methodologies to the down-sampled simulation data. Finally, some methods of refining the down-sampled data are applied and their effects are discussed.1. IntroductionSelective Laser Melting (SLM) is a common metal additive manufacturing process.

Finally, some methods of refining the down-sampled data are applied and their effects are discussed.1. IntroductionSelective Laser Melting (SLM) is a common metal additive manufacturing process. The process parameters determine the microstructure and, thus, the engineering properties. Infrared (IR) imaging is a non-contact detection method which is feasible for the in-situ monitoring of SLM processes [1]. The authors concluded that the thermal camera was able to capture the melt pool dimensions. However, the errors caused by quantization (digitization) in the binarized image was not considered (Fig. 1 (c)). They also related the lengths and widths of the melt pools with the laser scanning speeds, while these metrics are only measured from the binarized images [2]. Regarding the low-resolution temperature profile, they interpolated points into it. Wegner and Witt [3] used a thermal camera in the laser sintering process.

Regarding the low-resolution temperature profile, they interpolated points into it. Wegner and Witt [3] used a thermal camera in the laser sintering process. They used the melt’s temperature as an indicator of process parameters and related it with laser exposure time and laser power. An IR camera was shown to be also effective in locating defects during the SLM process of 6H-SiC [4].Work performed and funded by The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing & Technologies, LLC under contract number DE-NA0002839.173Fig. 1: Example of melt pool monitoring in an SLM process using a thermal camera: (a) thermal contour, (b) radiant temperature profile along scanning direction and (c) melt pool. [1]There are two types of thermal cameras: scanning devices and infrared Focal Plane Arrays (FPAs).

[1]There are two types of thermal cameras: scanning devices and infrared Focal Plane Arrays (FPAs). The FLIR thermal camera in the experimental platform at Missouri S&T is of the latter type and has a 640×512 pixel focal plane array with a 50 megapixel clock rate allowing full frame sampling at 126 frames/sec with a 14 bit dynamic range. The photodetectors in the FLIR thermal camera are the InGaAs array, which operate in the 0.9-1.7 μm wavelength region [5].The spatial resolution is an important parameter of the thermal camera. Each PhotoSensitive Element (PSE) in the array corresponds to one image pixel, The PSE signal can be expressed in volts, ADC units, the number of accumulated electrons, or the amount of stored charges. Its sources are always a photo current and a PSE dark current accumulated in the corresponding multiplexer capacitor [6]. In practice, the optical system can be configured to obtain a field of view at a distance from the build plate.

In practice, the optical system can be configured to obtain a field of view at a distance from the build plate. By configuring the optical system, a higher resolution can be achieved at the expense of the range of the field of view. When observing a large part or several parts together, the resolutions of the thermal images are usually very low. A trade-off must be made between the resolution and the viewing area size. Very-Low-Resolution (VLR) problems exist in image recognition, such as very-low-resolution human face recognition [7,8].A thermal camera has a specific resolution for a viewing area and the digital pictures are spatially discrete. The digital data can be treated as discrete samples of the continuously varying surface temperature distribution and each pixel in the picture is a single temperature measurement at its corresponding area. However, some of the spatial information will be lost by this down-sampling.

However, some of the spatial information will be lost by this down-sampling. In addition, the camera sensors collect the radiance inside the regions of the pixels, instead of at specific points. However, researchers often treat thermal images as discrete samples at specific grid points of surface temperature distribution and measure the melt pool’s174length and width directly from the image. The errors caused by low resolution sampling are not considered sometimes.In order to understand the effect of sampling, high-fidelity simulation results are utilized for investigation [9]. A sampling model is proposed and applied to the simulation data, and then artificial thermal images are generated. Features are measured from down-sampled images and original images separately, and their differences are just the errors caused by sampling. In other words, the measurements of the features are not exactly the real values.

In other words, the measurements of the features are not exactly the real values. The down-sampling has two main effects on the feature extraction: darkening effect of the melting pools and digitization of the morphological features. The darkening effect is caused by averaging the pixels and, thus, the maximal pixel value is always smaller than the true maximal surface temperature value.So some questions are brought up: how to understand and correct the measurements and recover the real features? In Section 2, a sampling model is proposed for the purpose of obtaining artificial low-resolution images from the original data, just like how the thermal camera works. In Section 3, the errors of the measurements of the features are plotted in figures, which shows the accuracy corresponding to the sampling grid size. In the following sections, interpolation and fitting methods are applied to the down-sampled data to find whether the methods improve the measurements.2.

In the following sections, interpolation and fitting methods are applied to the down-sampled data to find whether the methods improve the measurements.2. Sampling Model and Temperature Measurement ModelSince it is difficult to know the real surface temperature during the process, high-fidelity simulation data [9] of surface temperature are utilized instead of the real experimental surface temperature. The simulation data should be down-sampled to generate artificial thermal images with low resolutions. Therefore a sampling model is needed.The thermal camera collects the radiance inside the pixels over a period of time.

Therefore a sampling model is needed.The thermal camera collects the radiance inside the pixels over a period of time. Assumingthe emissivity ε of the material is constant, the sampled radiance is given by , tyxR , iiikttt x ix ix  2 x  2yyiy   2 y  2 , tyxT ,4 dxdydt,(1)x  2y  2x  2 , k is the gain and R is the radiance collected by the thermal camera sensor.y  2      y,ywhere T is the temperature inside the pixelx i,x iiit , t itime  ti  during theAssuming that the temperature field T is constant over the pixel area, (1) becomes  tyx 4RkT, tyx ,.(2)Solving for the measured temperature in equation (2) yieldsTM  R tyxk 4/1   ,(3)175where TM (K) is the temperature measurement. To recover the temperature from the measured radiance, one must determine the model parameters k and ε.

To recover the temperature from the measured radiance, one must determine the model parameters k and ε. Combining (1) and (3), the measurement function isT M   1 tyx ttt x  2 x  2x ixiyyy   tyxT , , 2 y  2i4 dxdydt4/1    .3. Errors from Spatial Down-SamplingThe features of a melt pool, namely the peak value p and melt pool length L, width W and area A, are measured from the thermal camera in practice, and the measurements of them have errors caused by low-resolution sampling. To determine the magnitude and nature of low-resolution sampling on these critical features, this section utilizes high-fidelity simulation data and the sampling model (4).3.1. Utilization of Simulation Data Simulation data of the surface temperature in the SLM process are utilized to characterize the errors resulting from the camera’s spatial resolution.

Utilization of Simulation Data Simulation data of the surface temperature in the SLM process are utilized to characterize the errors resulting from the camera’s spatial resolution. Using the down-sampling model, low-resolution images are obtained from the simulation data, which are used to investigate the errors caused by down-sampling. A piece of simulation video data is utilized for the investigation of sampling. In the simulation, the element size is Δx = Δy =10 μm and time step is Δt = 1 μs. The laser power is set to be 200 W and the laser pulse period is 90 μs. During each laser pulse period, the exposure time is 75 μs, and the laser is off for the 15 μs. The laser point distance is 60 μm. These manufacturing process parameters are often used when fabricating parts with 304L stainless steel [10]. The top and side views of a simulation frame are given in Fig. 2 and Fig. 3 respectively.Fig. 2: One Frame of Simulation176(4)Fig.

The top and side views of a simulation frame are given in Fig. 2 and Fig. 3 respectively.Fig. 2: One Frame of Simulation176(4)Fig. 3: Surface Temperature of a Simulation Frame (Lateral View)In this discrete case, where only spatial sampling is considered, equation (4) can be writtenas iT , M ， rssj , , 1 21     2 r rj jrm   11ri irn  11 4snsmT  1,24/1   ,where T(m, n) denotes the value at (m, n) in the original frame and TM(i, j, s1, s2, r) denotes the value at (i, j) in the down-sampled frame with camera grid shifting (s1, s2), and r is the down-sampling rate. In addition to the lower sampling rate r, it is important to note that the location of the sampling grid over the melt pool can have a critical impact on the measurements and measurement interpretation.

For example, a single pixel capturing the laser spot will show a significantly higher temperature than, for instance, if the laser spot fell at the intersection of four pixels. The parameters (s1, s2) are used here to capture melt pool locating effects within the camera pixels. To illustrate this effect with our simulation data (already sampled), in Fig. 4 the black lattice is the original image. In this case, the down-sampling rate is r = 5. The shifting of the blue lattice is (s1 = 1, s2 = 2) and that of the red lattice is (s1 = 2, s2 = 3). While it may be possible to carefully align the camera or reconstruct the precise laser path in order to determine the precise location of the laser spot at sub-pixel accuracy, here we assume that is not possible. Therefore, we consider such the location an uncertainty and evaluate the results of our low-resolution sampling with respect to this uncertainty.Fig. 4: Illustration of Sampling Grid Shifting177(5)3.2.

4: Illustration of Sampling Grid Shifting177(5)3.2. Direct Measuring Method Direct measurements are obtained by directly extracting the features from the thermal and the . The true peak temperature from thenmT ,images without data processing. Each frame is a matrix of temperatures down-sampled matrix of temperatures is simulation is iTMr,, sj 1,s,2pmax , nmnmT ,,and the peak temperature in the low-resolution image is  iT M s 1p,s,rmax , j iM2,, ssj , 12,r,The error caused by down-sampling with camera shift (s1, s2) and sampling rate r is , rtse,pM , rts , p .Measurement of the length, width and area of the melt pool are typically achieved by first defining a melt thresholds. A threshold is set, the matrix is searched, and all of the pixels greater than the threshold are identified.

A threshold is set, the matrix is searched, and all of the pixels greater than the threshold are identified. The region above the threshold is the set   , yxT i .E, yx i:thresholdiiThe length of the melt pool isLmax  x imin i x.Similarly, the width of the melt pool isWmax  yimin i y.Suppose N is the number of points in the set E and Ap is the size of each pixel, then the melt pool area isA pNA.In the down-sampled images, the measurements are defined as above and the measured length, width, and area, respectively, are denoted by LM, WM and AM. Then the length, width, area errors, respectively, aree LL ML,WWe W M,eAA MA.3.3. Errors from Down-Sampling Down-sampling leads to two notable effects on the measurements: the darkening effect on the peak temperature (Fig. 5) and the digitization of the morphological features (Fig. 6).

5) and the digitization of the morphological features (Fig. 6). The first effect is that the peak temperature in the low-resolution image is always lower than the original178(6)(7)(8)(9)(10)(11)(12)(13)(14)(15)peak temperature. The latter effect is that only several numbers can be taken as the measurements of the melt pool’s length, width and area from the low-resolution image.(a) Original Peak Temperature = 1924 K(b) Down-sampled Peak Temperature=1864 KFig. 5: Illustration of the Darkening Effect on the Peak Temperature(a) Original Melt Pool(b) Down-sampled Melt PoolFig. 6: Illustration of the Digitization of the Morphological FeaturesTo analyze the errors due to down-sampling, the data during one laser period are analyzed. In one experimental platform at Missouri S&T, a Renishaw system is fitted with a FLIR camera directed toward a portion of the powder bed with a 140 μm camera pixel size.

In one experimental platform at Missouri S&T, a Renishaw system is fitted with a FLIR camera directed toward a portion of the powder bed with a 140 μm camera pixel size. We will use this value as a nominal value for analysis and therefore set the down-sampling rate at r = 14. In the discrete simulation data, there are r2 = 196 total possible shifts in all. In this way, the distribution of the errors of the feature measurements are plotted in Fig. 7-Fig. 10. When defining the melt pool, take the threshold = 1700 K , which is the melt temperature of the material.In Fig. 7, it is shown that the measurements of the peak temperature will always be lower than the real value and the range of the peak temperature measurements is continuous and bounded. However, as for the morphological features (length, width, area), the measurements are highly digitized (Fig. 8-Fig. 10). In Fig. 8 and Fig. 9, the measurements of lengths and widths are in discontinuous groups.

8-Fig. 10). In Fig. 8 and Fig. 9, the measurements of lengths and widths are in discontinuous groups. The distance between the discontinuous bars is just the size the sampling grid, which is 140 μm in this case. In other words, the measurements can only be from {0, Δx, 2Δx, 3Δx, ...} if the camera pixel size is Δx. The digitization phenomenon happens on the area measurements similarly, where the distance between the bars in Fig. 10 is Δx2.179Fig. 7: Distribution of Peak Temperature ErrorsFig. 8: Distribution of Length ErrorsFig. 9: Distribution of Width ErrorsFig. 10: Distribution of Area ErrorsThe effect of spatial down-sampling may be instructive for the users who monitor the SLM process by a thermal camera. When choosing a size of the sampling grid and making all possible shifts, it is meaningful to calculate the maximum, the minimum and the mean of the errors caused by sampling with all shifts.

When choosing a size of the sampling grid and making all possible shifts, it is meaningful to calculate the maximum, the minimum and the mean of the errors caused by sampling with all shifts. The maximum, minimum and mean errors for the peak temperature and morphological features are plotted with respect to sampling grid sizes (Fig. 11-Fig. 14). These graphs show what resolution is need in order to achieve a desired level of accuracy. At a given sampling grid size (camera resolution), the maximum and the minimum of the measurements represent the uncertainty caused by the motion of the melt pool across the camera view, when one is conducting such an experiment.180Fig. 11: Peak Temperature Errors for Different Sampling GridsFig. 12: Melt Pool Length Errors for Different Sampling GridsFig. 13: Melt Pool Width Errors for Different Sampling GridsFig.

11: Peak Temperature Errors for Different Sampling GridsFig. 12: Melt Pool Length Errors for Different Sampling GridsFig. 13: Melt Pool Width Errors for Different Sampling GridsFig. 14: Melt Pool Area Errors for Different Sampling GridsThe results from the down-sampled simulation data show that the curves of peak temperature measurements with respect to the sampling grid sizes are smoother than the morphological features. Besides, when the spatial resolution is too low, the melting region cannot be extracted after down-sampling.4. Interpolation of Down-Sampled ImagesInterpolation is commonly used in image processing of low-resolution images. In this work, linear interpolation and spline interpolation are tried. The two interpolation methods have different natures. Linear interpolation is a simple method to create more pixels in a low-resolution image, but this method cannot produce a smooth curve. They are illustrated in Fig. 15 and Fig. 16.

Linear interpolation is a simple method to create more pixels in a low-resolution image, but this method cannot produce a smooth curve. They are illustrated in Fig. 15 and Fig. 16. Spline interpolation is a fitting method which uses a piecewise 3rd-degree polynomial function.181Fig. 15: Linear InterpolationFig. 16: Spline InterpolationHere the 2-dimensional interpolations are applied to the artificial down-sampled images, and the interpolation rate should be such as to restore the original resolution. The following are the results of measurements after linear interpolation (Fig. 17-Fig. 20).Fig. 17: Peak Temperature Errors after Linear InterpolationFig. 18: Melt Pool Length Errors after Linear InterpolationFig. 19: Melt Pool Width Errors after Linear InterpolationFig. 20: Melt Pool Area Errors after Linear Interpolation182The following are the results of measurements after spline interpolation.Fig. 21:Peak Temperature Errors after Spline InterpolationFig.

20: Melt Pool Area Errors after Linear Interpolation182The following are the results of measurements after spline interpolation.Fig. 21:Peak Temperature Errors after Spline InterpolationFig. 22: Melt Pool Length Errors after Spline InterpolationFig. 23: Melt Pool Width Errors after Spline InterpolationFig. 24: Melt Pool Area Errors after Spline InterpolationThe simple linear interpolation can reduce the digitization effect a little. But linear interpolation cannot restore the peak value which is lost by sampling, just as illustrated in Fig. 15. The spline interpolation is hopeful to restore the peak value of a curve (Fig. 16). However, it does not make obvious improvements in the measurements of features, but makes the measurements worse sometimes. The reason is that interpolation methods treat every pixel point as a pointwise measurement on the original surface, but in fact the pixel value is a sort of average over the pixel region.

The reason is that interpolation methods treat every pixel point as a pointwise measurement on the original surface, but in fact the pixel value is a sort of average over the pixel region. So pointwise interpolations are not proper in the problem.5. Fitting MethodAn alternative to direct measurement from thermal camera data is to fit the measurements to a model or set of basis functions. Here, we explore a fitting method that segments the melt pool into two parts: the front part and the rear part, as illustrated in Fig. 25. The temperature of the front part is simpler and can be approximated by a Gaussian-like function183 , yxTAexp   xx  0 2 S 12y Sy 2 202   Tamb,xx 0,where T(x, y) is the surface temperature, (x0, y0) is the peak temperature location and Tamb = 353 K is the ambient temperature in the simulation.

For simplicity, let (x0, y0) = (0, 0) by coordinate system translation, and then (16) can be written as , yxTAexp  2x 2 S 1y S22 2  Tamb,x0,The down-sampled image is shown in Fig. 26, where the sampling grid size Δx = Δy =40 μm. The peak temperature after down-sampling (pM = 2233 K) is lower than the original peak temperature (p = 2371 K).A fitting method is proposed to have a better estimation of the peak temperature based onthe low-resolution image. The penalty function is , xSSAP 1 0,,2,y0         i4, xSSAtsT 0,,,,12,y0 dtds4/1    iT   2,2/ x ix,2/x ix 2/y  iyy,2/and (xi, yi, Ti) are the data set for y where i fitting. The optimization is fulfilled by “fmincon” function in Matlab. The estimated peak temperature from the fitting result is pest = 2338 K.

The optimization is fulfilled by “fmincon” function in Matlab. The estimated peak temperature from the fitting result is pest = 2338 K. In this way, the estimated peak temperature is better than that from the low-resolution image.iFig. 25: Surface Temperature in SimulationFig. 26: Down-Sampled Surface Temperature184(16)(17)(18)Fig. 27: Fitting Result of the Front Part of the Melt PoolUsing this method, the width of the melt pool can also be estimated. The original melt pool width is W = 120 μm. The down-sampled melt pool width is WM = 80 μm. The estimated width is West = 112 μm. The fitting method can also obtain a better measurement of the melt pool width. The fitting method is based on the assumed surface temperature function. The Gaussian-like function (17) is not applicable to the rear part of the melt pool, and the existence of the phase transition region is still a challenge for the fitting method.6.

The Gaussian-like function (17) is not applicable to the rear part of the melt pool, and the existence of the phase transition region is still a challenge for the fitting method.6. ConclusionsThe thermal camera is an instrument for monitoring the SLM process. Temperature information and morphological information of melt pools are often related to manufacturing parameters and engineering properties. It is worth noting that the features extracted from low-resolution thermal images contain errors caused by spatial sampling. In this work, high-fidelity simulation data are utilized and a sampling model is proposed. Based on these, the errors caused by sampling grid sizes are plotted for all possible grid shifts. This gives the level of accuracy for a given camera pixel size, which is worth of concern in experiments. Interpolation is often used in the processing of low-resolution images, but interpolation is not enough to eliminate the effect of sampling.

Interpolation is often used in the processing of low-resolution images, but interpolation is not enough to eliminate the effect of sampling. The model-based function fitting method can improve the measurements of the melt width and peak temperature.7. AcknowledgementThis work was funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.1858. References[1] Cheng, B., Lydon, J., Cooper, K., Cole, V., Northrop, P., & Chou, K. (2018). Infrared thermal imaging for melt pool analysis in SLM: a feasibility investigation. Virtual and Physical Prototyping, 13(1), 8-13.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEffects of Unit Cell Size on the Mechanical Performance of Additive Manufactured Lattice StructuresArash SoltaniTehrani1, 2,Seungjong Lee1, 2,MohammadReza VaziriSereshk ,Nima Shamsaei1,2, *1National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849 2Department of Mechanical Engineering, Auburn University, Auburn, AL 36849Corresponding author: shamsaei@auburn.eduAbstractLattice structures are generated through the repetition of smaller structures, defined as unit cells. These structures are popular alternatives for bone implants due to the potential to adjust the stiffness. However, in some applications, there are volume and mass constraints that cannot be exceeded.

These structures are popular alternatives for bone implants due to the potential to adjust the stiffness. However, in some applications, there are volume and mass constraints that cannot be exceeded. Therefore, to match the lattice structure’s stiffness to that of the natural bone, unit cell sizes should be altered. In this study, the effects of different unit cell sizes, on the compression behavior of lattice structures fabricated from 316L stainless steel (SS) via laser beam powder bed fusion (LB-PBF) are studied through finite element analysis (FEA) while the volume and mass are kept constant and results of which, are validated by experiments. It was found that energy absorption capability and stiffness of lattice structures can increase with decreasing the size while the volume and mass are kept constant.

It was found that energy absorption capability and stiffness of lattice structures can increase with decreasing the size while the volume and mass are kept constant. The lattice structure with smaller unit cell dimensions tolerated a relatively higher maximum force for the same amount of displacement.Keywords: Lattice Structures; Laser Beam Powder Bed Fusion (LB-PBF); Stainless Steels;Stiffness; Bone ImplantsIntroductionLattice structures are porous structures which are generated through the repetition of smaller structures defined as unit cell [1]. Unit cells are also formed by the connection of struts with different cross-section shapes such as rectangular, circular and hexagonal [1]. Some common unit cells are BCC, FCC, diamond, and gyroid [1–4].

Some common unit cells are BCC, FCC, diamond, and gyroid [1–4]. Therefore, different mechanical properties can be expected by the different unit cell types, which makes lattice structures suitable for many applications such as medical implants, energy absorption, and light-weighting applications [1,5– 9]. However, fabrication of lattice structures via traditional manufacturing methods (i.e., cutting, machining, brazing, and welding) is difficult or even impossible.Nowadays, producing lattice structures with different additive manufacturing (AM) technologies such as laser beam powder bed fusion (LB-PBF) is very common due to more design freedom in these manufacturing methods. In LB-PBF, parts are fabricated in a layer-wise fashion from the powder feedstock with minimal interference of human in manufacturing. Different computer-aided design (CAD) files can be imported directly to the manufacturing machine for production.

Different computer-aided design (CAD) files can be imported directly to the manufacturing machine for production. In addition, processing many materials such as titanium alloys (e.g., Ti-6Al-4V),2254aluminum alloys (e.g., AlSi10Mg), and stainless steel (SS) alloys (e.g., 316L SS) [1–6,9–14] has been made possible through this type of manufacturing.Although manufacturing lattice structures with very complex geometries and from different materials have been possible via LB-PBF, there are still some challenges, which have hampered their widespread use in different applications. For example, when the lattice structure is used as a bone implant, stress shielding can occur due to the stiffness mismatch between the bone and implant [15]. Fatigue failure can also happen in implants since they are usually exposed to different loading types such as compression-compression and compression-tension [1].

Fatigue failure can also happen in implants since they are usually exposed to different loading types such as compression-compression and compression-tension [1]. The relatively high surface roughness after fabrication via LB-PBF can also be detrimental in some applications [16]. These challenges make studying different lattice structures very important. Many studies have been performed regarding lattice structures fabricated with different unit cell types. For example, Yanez et al. [14] studied gyroid lattice structures fabricated via electron beam melting (EBM) in normal- and deformed-shape conditions. Higher stiffness and strength were observed for deformed gyroids under compression while normal gyroids showed higher stiffness and strength under torsion.Van Hooreweder et al. [17] studied different post-processing treatments including hot isostatic pressing (HIP) and chemical etching in order to decrease the surface roughness inherent to LB-PBF.

[17] studied different post-processing treatments including hot isostatic pressing (HIP) and chemical etching in order to decrease the surface roughness inherent to LB-PBF. Chemical etching was observed to increase fatigue strength of Ti-6Al-4V samples. In addition, further improve in fatigue strength was observed by HIP specially in the high-cycle- fatigue (HCF) region. Lieteart et al. [18] studied different loading combinations including tension- tension, compression-tension, and compression-compression to investigate the fatigue behavior of lattice structures. It was observed that due to the presence of local mean tensile stress in tension- tension and compression-compression loading types, lower fatigue life can be yielded compared with compression-tension loading. Plessis et al. [12] showed that different unit cell types can also result in different mechanical properties while the volume fraction is constant.

Plessis et al. [12] showed that different unit cell types can also result in different mechanical properties while the volume fraction is constant. In this study, diagonal and rhombic lattice structures fabricated from Ti-6Al-4V via LB-PBF with a 50% volume fraction under compressive loading were investigated [12]. Relatively higher ultimate compressive stress (UCS) and Young’s modulus were observed for the rhombic lattice structure compared with diagonal. μCT simulations also revealed that failures started from high-stress regions which were mostly located at the nodes.In addition to the aforementioned challenge, there are sometimes mass and volume constraints as well for specific applications. In order to obtain the desired mechanical properties (e.g., stiffness) and keep constant mass and volume, unit cell dimensions need to be altered.

In order to obtain the desired mechanical properties (e.g., stiffness) and keep constant mass and volume, unit cell dimensions need to be altered. In this study, effects of unit cell dimensions on the mechanical behavior of lattice structures under static loadings will be investigated via finite element analysis (FEA). The results of FEA will also be verified by experimental quasi-static compression tests, and final conclusions will be made based on the results.Finite Element SimulationsLattice structures were analyzed with ABAQUS in order to study the mechanical response under monotonic compression. A linear tetrahedral element (C3D4) was used to mesh the structures. The mesh was also verified to rule out the mesh-dependency of the results. Both elastic and plastic properties were selected for the material according to 316L SS properties obtained from [19].

The mesh was also verified to rule out the mesh-dependency of the results. Both elastic and plastic properties were selected for the material according to 316L SS properties obtained from [19]. A 5-mm displacement and an ENCASTRE boundary conditions were applied on the top and2255bottom plates of lattice structures, respectively. Von Misses stress, displacement, reaction forces, and strain were obtained using FEA in lattice structures.Experimental ProgramDifferent lattice structures were fabricated from different BCC4 unit cells via Renishaw AM250, an LB-PBF manufacturing machine, in an argon atmosphere (see Figure 1(a)). The powder feedstock used in this study was argon-atomized 316L SS fabricated by LPW Technology Ltd with an initial particle size distribution of 15-45 μm. The overall volume of lattice structures was designed with dimensions of 14.4 mm × 14.4 mm × 14.4 mm.

The overall volume of lattice structures was designed with dimensions of 14.4 mm × 14.4 mm × 14.4 mm. Lattice structures were fabricated with 6, 8, and 10 unit cells in each direction with different unit cell lengths and strut diameters in order to keep volume and mass constant. Two plates with a thickness of 0.5 mm were also located on top and bottom of lattice structures to uniformly distribute the compressive load. As shown in Figure 1(b), lattice structures with 6, 8, and 10 unit cells in each direction were defined as BCC6, BCC8 and BCC10, respectively.JCXXXXXXXXX>II • X:x: X)C)CJCJCX: X)C)C)C)CJC~'!Ill)C)C)CI BCC6 II Bees II BCClO IFigure 1 (a) isometric (left) and top (right) view of the build plate, and (b) front view of BCC6, BCC8, andBCC10After fabrication, specimens were removed from the build plate.

They were cleaned completely using an ultrasonic cleaner to ensure removal of loose powder particles entrapped within the lattice structures or on the surface. Parts were measured and tested without any additional post-processing treatment. Quasi-static compression tests were performed via MTS 810 servo-hydraulic system with a displacement rate of 0.015 mm/s. Tests were performed in a displacement-controlled mode and continued to 5 mm. The tests were stopped if the applied force2256reached 70 kN and considered as failed. To verify the results, three tests were performed for each condition.Results and discussionLattice structures including BCC6, BCC8, and BCC10 were analyzed using finite element method. All lattice structures went through 5 mm displacement (see Figure 2). As seen in this figure, the BCC6 lattice structure under monotonic compression deforms with a barreling shape.

All lattice structures went through 5 mm displacement (see Figure 2). As seen in this figure, the BCC6 lattice structure under monotonic compression deforms with a barreling shape. In addition, the first row of unit cells under compression have the highest amount of displacement leading to collapse of the first row prior to other rows. This behavior continues in the aforementioned lattice structure until all the rows collapse and there is no unit cell to absorb the energy. In this stage, densification in lattice structure happens which comes with a sudden increase in stress and final failure of the specimen [1]. Similar behavior in displacement from FEA was observed for BCC8 and BCC10. However, it was observed that BCC8 and BCC10 can absorb more energy compared with BCC6 lattice structure. It seems that the final failure in BCC10 happens at a relatively higher strain compared with BCC8 and finally BCC6.

It seems that the final failure in BCC10 happens at a relatively higher strain compared with BCC8 and finally BCC6. Therefore, it is expected that for the same amount of displacement, BCC10 requires more applied force compared with BCC8 and BCC6, respectively. This behavior can be explained by the existence of more unit cells to absorb the applied energy.,: /.. I· -~✓/ \_,fI' 1? "'- "'-/,fv¥! . ,,,, /,f'\,✓,;,'1 o/f .~m, • " / : , I~ I i f -~ ~ r" "'''·""...,.,f ~·~Figure 2 BCC6 lattice structure in (a) before and (b) after 5 mm displacementAs seen in Figure 3, Von Misses stress within the BCC6, 8, and 10 lattice structures is shown for 5 mm displacement. No failure was observed by deforming the lattice structure for 5 mm. In addition, it can be observed that a relatively smaller amount of stress is applied on the vertical struts of BCC10 followed by BCC8 compared with BCC6. This behavior was also verified by looking at the strain results.

This behavior was also verified by looking at the strain results. It was observed that the amount of strain at the vertical struts of BCC6 lattice structure is at the highest. This decrease in stress and strain was also observed for the internal diagonal struts of each unit cells, however, it was more significant for the vertical struts. Moreover, as seen in this figure, the horizontal struts of all lattice structures are under tension due2257..,to bending stress at these struts although the structure is under compression. Therefore, horizontal struts under relatively higher tension deform more in the perpendicular direction to loading, which contributes to the higher energy absorption of BCC10 lattice structure followed by BCC8 and BCC6.The reaction forces from the top plate were also plotted for deformed lattice structures using FEA, and it was observed that the maximum reaction force in BCC6, BCC8, and BCC10 are 539 kN, 402 kN, and 280 kN, respectively.

However, there were relatively higher number of regions with maximum force in the BCC10. This shows that by increasing the number of unit cells, the applied force, and consequently, stress is divided into more unit cells which decrease the maximum amount of stress. Therefore, smaller displacement happens at BCC10 unit cells for the same amount of applied stress.Figure 3 Von Misses stress shown in (a) BCC6, (b) BCC8, and (c) BCC10 lattice structures after 5 mmdisplacementIn addition to numerical simulations, all lattice structures with different number of unit cells in each direction were tested under monotonic compressive loading. The force-displacement curves are shown in Figure 4. As shown in this figure, all lattice structures are first deformed with a linear elastic behavior, which is then followed by a plateau region where force gradually increases with permanent deformation. This behavior conforms with the literature [20].

This behavior conforms with the literature [20]. In these studies, the third regime (i.e., densification) where a sudden increase in force/stress happens for a relatively small increment of displacement, is also shown. However, in the current study, tests were stopped at 5 mm, and the third regime was not shown. It can be observed from the figure that BCC8 and BCC10 were almost comparable in the elastic region. However, BCC10 showed superior strength after the first yield.2258403530,-._ ~ 25 ' - ' ~ 20 u ;... 0 µ.. 15105BCC6 - BCC8 - BCClO00 0.511.52 2.53 3.5 4 4.55Displacement (mm) Figure 4 Force-displacement curves for BCC6, BCC8 and BCC10 lattice structures under displacement-controlled compressive loading for 5 mm deformationAll lattice structures deformed comparable to each other, and deformation started with the failure of unit cells row by row.

In addition, all the specimens deformed with a barreling shape, which was also observed in FEA results (see Figure 5). Stiffness was calculated based on the slope of the force-displacement curve in the elastic region for each lattice structure, and it was found to be 18.4 N/mm, 22.9 N/mm, and 24.9 N/mm for BCC6, BCC8, and BCC10, respectively. Therefore, stiffness was increased by increasing the number of unit cells and decreasing the dimensions (i.e., unit cell length and strut diameter). Since in open-celled structures (i.e., lattice), bending deflection is proportionate to force, length, Young’s modulus, and moment of inertia [21], it can be concluded that stiffness is directly correlated with Young’s modulus and moment of inertia, and inversely correlated with length.

Therefore, by increasing the number of unit cells, and consequently, decreasing the length of unit cells, an increased stiffness value can be expected.In all deformed lattice structures, deformation started from the nodes between each two rows. This behavior conforms well with the FEA where a relatively higher Von Misses stress was shown at these nodes. It was observed that stress is relatively higher in vertical struts of BCC6 while by increasing number of unit cells, a smaller amount of stress was applied on these struts. Although the lattice structure is tested under compressive loading, there was also tensile stress available in the horizontal struts due to the bending moment.

Although the lattice structure is tested under compressive loading, there was also tensile stress available in the horizontal struts due to the bending moment. Therefore, unit cells fail at nodes where horizontal and vertical struts are connected to each other, and this behavior continues until the part is completely failed.22598illFigure 5 BCC10 lattice structure: (a) before and (b) after 5 mm displacementConclusionsIn this study, three different lattice structures fabricated from BCC4 unit cell with different dimensions (i.e., unit cell length and strut diameter), were fabricated via LB-PBF from 316L SS while the volume and mass were kept constant. Finite element analysis was used to investigate the behavior of these structures under compression and to find the locations for maximum stress. In addition, monotonic behavior of these structures was obtained under compressive loading with a displacement-controlled program, where deformation was set to be 5 mm.

In addition, monotonic behavior of these structures was obtained under compressive loading with a displacement-controlled program, where deformation was set to be 5 mm. Based on the observations and results, following conclusions can be made:1. BCC4 lattice structures represented good strength and stiffness, which make them suitable for many medical applications [22].2. The monotonic behavior of lattice structures under compressive loadings did not alter significantly with the change of unit cell dimensions, while the weight and mass were kept constant. It was observed that stiffness of the structure increases with increasing the number of unit cells as well as decreasing the unit cell sizes.3.4. Decreasing the number of unit cells while the volume and mass were kept constant resulted in higher energy absorption and less deformation under similar amount of applied force. References[1]X.Z. Zhang, M. Leary, H.P. Tang, T. Song, M.

E EFFECTS OF 3D PRINTED MOLDS ON METAL CASTINGSDean Snelling1, Heather Blount2, Charles Forman2, Kelly Ramsburg2, Andrew Wentzel2, Christopher Williams1, Alan Druschitz2 1 Design, Research, and Education for Additive Manufacturing Systems Laboratory; Department of Mechanical Engineering 2 Department of Material Science Virginia Tech, Blacksburg, VAABSTRACTAdditive manufacture of sand molds via binder jetting enables the casting of complex metal geometries. Various material systems have been created for 3D printing of sand molds; however, a formal study of the materials’ effects on cast products has not yet been conducted. In this paper the authors investigate potential differences in material properties (microstructure, porosity, mechanical strength) of A356 – T6 castings resulting from two different commercially available 3D printing media.

In addition, the material properties of cast products from traditional “no-bake” silica sand is used as a basis for comparison of castings produced by the 3D printed molds.Keywords: Binder Jetting, Indirect 3D Printing, Metal Casting, Sand Casting1.EFFECTS OF MOLDING MATERIALS ON CASTINGS1.1.Additive Manufacturing of Sand Molds for Metal CastingAdditive Manufacturing (AM) has enabled the direct production of molds without the need for a pattern. Specifically, the binder jetting 3D Printing (3DP) process has been used to directly fabricate sand molds and core boxes by selectively jetting binder into a powder bed of foundry sand [1].A schematic of the binder jetting process is provided in Figure 1. During binder jetting, a polymer binder is printed onto a bed of powder using a traditional inkjet print head to form one cross-sectional layer of the part. After a layer of binder is printed, the powder bed lowers, and fresh powder is spread over the powder bed using a roller.

After a layer of binder is printed, the powder bed lowers, and fresh powder is spread over the powder bed using a roller. Then, the next layer of the part is patterned onto the powder bed atop the previous layer. In this manner, the object is constructed layer by layer.After the mold is printed, the excess powder is removed using compressed air or vacuum. Often, the printed molds are then cured in an oven to eliminate a portion of the binder (Figure 2a). The printed part can also be used as a core or as a complete mold, which includes runners, gates, and a down-sprue (Figure 2b) [2]. Molten metal is then cast into the mold to create thecasting (Figure 2c). 3DP of molds for metal casting has been commercialized by a variety of companies, including 3D Systems [3], ExOne [4], and Viridis3D [5].Figure 1. Schematic of the binder jetting processFigure 2.

3DP of molds for metal casting has been commercialized by a variety of companies, including 3D Systems [3], ExOne [4], and Viridis3D [5].Figure 1. Schematic of the binder jetting processFigure 2. a) Complex printed mold created using binder jetting, b) no-bake outer mold, and c) cast complex structureMany of the applications for additively manufactured sand molds are in providing a means to offer flexible tooling for traditionally designed castings. However, an important asset of the process is that the geometric freedoms offered by AM can be leveraged to provide a means for metal casting of complex geometries that are not possible to fabricate via traditional casting means [6]. In addition, the layer-by-layer process of fabricating sand molds enables a designer to uniquely integrate vents, sprues, runners directly into the mold design.

In addition, the layer-by-layer process of fabricating sand molds enables a designer to uniquely integrate vents, sprues, runners directly into the mold design. Finally, as the final part is created outside of the AM systems’ build chamber, the binder jetting and casting process chain enables the creation of large metal geometries. Specifically, multiple molds may be printed and fitted together with core paste to pour large metal castings.1.2.Traditional Molding Material Effects on CastingsAlthough AM of sand molds has enabled designers to overcome manufacturing restrictions, little is known about the materials systems’ effects on metal castings. This gap in knowledge is contradictory to the knowledge base in traditional sand casting.

This gap in knowledge is contradictory to the knowledge base in traditional sand casting. Many of the common aggregates used for the formation of molds in traditional sand casting are comprised of silica sand, natural minerals, synthetics, and other particulate materials [7]; each component with differing characteristics such as composition, grain size, and binder or compaction requirement. As a result, the properties of subsequent metal castings vary due to their reaction to the mold. For example, the quality of a casting can change due to water vapor stored in the mold, freehydrogen, and organics as the metal flows and solidifies [7].

For example, the quality of a casting can change due to water vapor stored in the mold, freehydrogen, and organics as the metal flows and solidifies [7]. These reactions inherently affect the final cast product and can produce defects such as porosity, oxidation, carbon formation, and surface roughness [7].1.3.Prior Research in 3DP of Sand MoldsThis existing understanding of sand molds on metal casting cannot be directly applied to AM sand materials systems as they are the working materials are dissimilar. There is limited literature that explores the effects of 3D printable mold material on castings. Instead, most of prior research is focused in studying the molds produced by ZCast® [8–10]. For example, McKenna et al. performed tests on ZCast®, to determine the effects of temperature and curing time on permeability and compressive strength of the mold.

For example, McKenna et al. performed tests on ZCast®, to determine the effects of temperature and curing time on permeability and compressive strength of the mold. A mathematical model was used to determine an optimal curing time and temperature for both permeability and compressive strength [11]. In previous work, the authors investigated the binder content of ZCast® material system and found it had a significantly higher binder content (up to 8% binder) than traditional no-bake foundry sand [2]. The higher binder content of the ZCast® printed molds causes molds to generate more gas during casting, which can cause defects in the final parts [2]. A new curing cycle with higher temperatures for a shorter duration produced more consistent cast structures with fewer gas defects [2].

A new curing cycle with higher temperatures for a shorter duration produced more consistent cast structures with fewer gas defects [2]. In addition, Gill & Kaplas compared castings printed with ZCast® and Investment casting using starch and plaster, including dimensional tolerances, hardness values, surface roughness, production cost, and shrinkage [12]. Experiments were also run at different shell thicknesses. It was determined that starch based investment casting produced higher hardness values and slightly better surface roughness, where ZCast produced better dimensional tolerances all from a recommended shell thickness range of 12-2mm [12]. It was also concluded there is optimal settings in terms of time and shell thickness to minimize cost based on individual builds [12].1.4. Overview of WorkIn order to ensure quality cast parts, the effects of final cast material properties needs to be studied using different molding materials.

Overview of WorkIn order to ensure quality cast parts, the effects of final cast material properties needs to be studied using different molding materials. The primary goal of this work is to compare two commercially available 3D printing 3DP powders: ViriCast™ (produced by Viridis3Da) and ZCast® (produced by 3D Systemsb). Additionally, this work compares these two 3DP powders on the basis of the handleability of the resultant printed molds and the properties of the cast metal parts they produce. The 3DP powders will also be compared to traditional no-bake foundry sand in order to determine whether 3D printed molds can produce metal parts of comparable quality to traditional casting approaches.An explanation of multiple tests utilized to characterize both 3DP sands and no-bake sand and material properties will be presented in Section 2. These tests include sieve analysis, tensile testing, surface roughness, density, hardness, porosity, microstructure, and compression tests.

These tests include sieve analysis, tensile testing, surface roughness, density, hardness, porosity, microstructure, and compression tests. The results of these tests are presented and statistically analyzed in Section 3 Finally, in conclusion, an overview and future work is given in Section 4.a http://www.viridis3d.com/metalcasting.htm b http://www.zcorp.com/en/Products/3D-Printers/Spectrum-Z510/spage.aspx2.EXPERIMENTAL PROCEDURE2.1.Characterization of 3D Printing PowderFunctionally, the two sands investigated in this paper are similar. They are both processed via a binder jetting AM process, and are both designed for receiving molten metal for sand casting applications. The only stated performance difference is in their maximum casting temperatures: ViriCast™ can be cast at 2650°F (1454.4 °C) and can be used to cast ferrous and non-ferrous alloys; ZCast® has a maximum cast temperature at 2000°F (1093.3 °C) and can only be used for casting non-ferrous alloys.

Further differences between the two powders are identified by further characterization experiments, including sieve analysis and tensile testing of cured sand samples.Sieve AnalysisStandard sieve tests according to AFS 1105-00-S [13] and AFS 1106-00-S [13] were previously performed on ZCast® powder and no-bake sand to determine particle size distribution [2]. A sieve test was performed on the ViriCast™ powder following the same standard procedure.Tensile TestingThe mechanical strength of the molds was characterized via tensile testing. Tensile testing was previously performed on ZCast® powder and no-bake foundry sand according to AFS 3342- 00-S to determine handleability of the material [2]. Five equivalent dog-bone shaped specimens were printed using a 3D Systems Spectrum Z510 3D printer with ViriCast™ powder and 3D Systems zb56 binder; the dog-bones were then cured at 400 °F (204.4 °C) for five hours [2].

Tensile testing was performed using a tensile testing machine to determine the mold fracture strength. Collected data and modes of fracture for ViriCast™ molds were compared with ZCast®, as well as no-bake foundry sand molds.2.2.Characterization of Cast A356 CylindersThe primary goal of this research was to compare the properties of metal cast in (i) ViriCast™ molds, (ii) ZCast® molds, and (iii) no-bake foundry sand molds. The two 3D printed sands are treated as the experimental group; chemically bonded silica sand, also known as no- bake sand, is treated as the control, since there exists published information about its casting properties [14]. The binder ratio in the no-bake sand was 4:1 of Phenoset RB to APR-015 hardener/catalyst, which accounted for approximately 1.6% of the sand mixture [2].

The binder ratio in the no-bake sand was 4:1 of Phenoset RB to APR-015 hardener/catalyst, which accounted for approximately 1.6% of the sand mixture [2]. A Palmer M50XLD continuous sand mixer was used to mix the silica sand and binder to create the no-bake sand [2].To make the 3DP molds, hollow cylinders (inner diameter one inch, wall thickness of 1 inch, and length of 4 inches) were designed using CAD. This specific inner diameter was chosen given the ability to machine to typical specimen sizes for ASTM compression tests given in ASTM E9-09 [15]. Six cylindrical molds where then printed in both ViriCast™ and ZCast® powders at their individual manufacturer’s process parameter specifications (Table 1).

Six cylindrical molds where then printed in both ViriCast™ and ZCast® powders at their individual manufacturer’s process parameter specifications (Table 1). The printed mold can is illustrated in Figure 3a and the manufacturer’s process parameters in Table 1.The resultant printed molds were then post-processed according to their manufacturer’s specifications: ViriCast™ molds were cured at 400°F (204.4°C) for five hours, and ZCast® molds were cured at 600 °F (316 °C) for one hour. Then, no-bake foundry sand was used to create the down-sprue, runners and gates. In order to create the no-bake molds, one inch diameter dowel rods were used to create four cylindrical molds in no-bake foundry sand.Table 1. 3D Printed mold material manufacturer process parameter specifications3D Printed MaterialSaturation LevelBinder/Volume RatioZCast®Shell Core94% 49%0.204517 0.0530748ViriCast™Shell Core85% 120%0.184935 0.129979A356 alloy was cast into all the molds.

A standard T6 heat treatment of 1005°F (540.6 °C) for six hours and artificial aging at 315°F (157.2 °C) for five hours was applied to the cylinders. The cylinders were cut into top, middle, and bottom sections for material analysis as shown in Figure 3b.Figure 3. a) 3D printed cylindrical mold and b) cast cylinder with diagram of cylinder sectionsTwo top sections, two middle sections, and two bottom sections of the cylinders cast from each mold material were analyzed for surface roughness, density, hardness, porosity, and microstructure. The remaining specimens were machined for compression testing.

The remaining specimens were machined for compression testing. Average values for the overall cylinders are presented along with standard deviation to aid analysis.Surface RoughnessSurface roughness was measured using a Phase II SRG Surface Roughness Tester.Roughness average (Ra) of the cylinders cast using different molds was measured.DensityArchimedes’ Principle was used to determine the density of the cast A356 aluminum cylinders created by each of the mold types. Three trials were conducted for each cylinder section.HardnessSamples were mounted in PhenoCure® Resin Powder and burnished to with 240, 320, 400, and 600 grit polishing paper using the Ecomet® 3 Variable Speed Grinder-Polisher. Hardness testing was performed using a LECO Vickers Hardness Tester LV700AT.

Hardness testing was performed using a LECO Vickers Hardness Tester LV700AT. The cross-section hardness was measured in five locations of each cylinder.PorosityThe sample surfaces were ground to remove indentations from hardness testing, and then re- polished using 240, 320, 400, and 600 grit polishing paper. Polishing was conducted using 5 (cid:1)(cid:2) and 1 (cid:1)(cid:2) alumina suspensions and a final finishing cloth. Final polishing was conducted with 0.04 (cid:1)(cid:2) colloidal silica and a final finishing cloth. Nine optical micrographs were taken of each sample for porosity measurements. ImageJ software was used to find the percent porosity by calculating the percentage of the total area covered by pores in each micrograph [16]. To accomplish this, the software was used to adjust the threshold of the image, highlight the pores, and measure the percent area of the pores.

To accomplish this, the software was used to adjust the threshold of the image, highlight the pores, and measure the percent area of the pores. The threshold color brightness was adjusted until the pores were fully highlighted, and the size settings for analyzing particles were adjusted until the software recognized the pores. The ImageJ settings depended on the original saturation and contrast of the images. For example, micrographs with less contrast between black pores and surrounding material require higher threshold color brightness settings in ImageJ to fully highlight pores. Table 2 shows the ImageJ settings used to calculate porosity.Table 2.

Table 2 shows the ImageJ settings used to calculate porosity.Table 2. ImageJ settings for calculating porosity of aluminum cylindersMold Material Threshold: Color Brightness Analyze Particles: Size 3DP Powders No-Bake28 95115-Infinity 15-InfinityMicrostructureNext, the microstructure of the aluminum samples were revealed by etching with Weck’s reagent [17], [18] , which contains 100 mL water, 4 g KMnO4, and 1 g NaOH. The sample surfaces were submerged in Weck’s reagent [17], [18] and agitated for 20 seconds. After rinsing with water and alcohol, the samples were blown dry. Optical microscopy was performed to characterize the microstructure and determine the dendrite arm spacing in each sample.Compression TestingCompression specimens were machined to a diameter of ½ in. and length of 1 in. according to ASTM standard E9-09 [15].

and length of 1 in. according to ASTM standard E9-09 [15]. Compression tests were conducted using an MTS Insight Electromechanical 150 kN Standard Length Testing System to measure the compressive yield strength. The strain rate was fixed at 0.005 in./min. Compressive yield strengths were found using a 0.02% offset from the elastic region of the stress-strain curve. Compression tests were not performed on the cylinders cast using no-bake foundry sand since the compression behavior cast T6-A356 aluminum is published information [14].3.RESULTS AND DISCUSSION3.1.Properties of 3D Printing PowderSieve AnalysisPrevious testing revealed that the no-bake sand had an AFS grain fineness number (GFN) of 57, while the ZCast® powder had an AFS GFN of 143 [2]. The results of the sieve analysis from silica sand and ZCast® are seen in Table 3 and ViriCast™ in Table 4.Table 3.

Regardless, the sieve analysis demonstrated that the ViriCast™ powder is significantly finer than the ZCast® powder, and both 3DP powders are much smaller in size than the no-bake sand.Tensile TestingDog bone shaped specimens were produced from the mold materials and tested following theexperimental plan (Section 2.1). The results of tensile testing are reported in Table 5.Table 5. Tensile strengths of the different mold materialsMold MaterialNo-Bake ViriCast™ ZCast®Mean σ (MPa) 0.56 0.16 0.06SD (MPa) 0.09 0.03 0.02Wilcoxon Comparison p-ValueX 0.0081 0.0369-------- X 0.0282Statistical software (JMP 10.0.2) was used to investigate potential significant difference in the tensile strengths of the two printed mold materials. Non-parametric Wilcoxon tests were performed on the data using a level of significance (α) of 0.05 after confirming the data was not normal. The p-values from the comparisons are displayed in Table 5.

The p-values from the comparisons are displayed in Table 5. The no-bake molds were significantly stronger than the two 3DP materials. Additionally, the ViriCast™ material was significantly stronger than the ZCast®. While ViriCast™ was weaker than the traditional no- bake foundry sand, it would make a more durable and handleable mold than the ZCast® 3DP powder.3.2.Properties of Cast A356 CylindersCylindrical specimens cast from the two 3DP mold materials were compared to sand casted A356-T6 alloy cylinders of the same size. The metal specimens cast in no-bake molds were prepared using a different A356 melt than the metal samples cast using 3DP molds. The same T6 heat treatment was performed on the metal specimens produced with both the 3DP and no- bake sand molds.

The same T6 heat treatment was performed on the metal specimens produced with both the 3DP and no- bake sand molds. Compositional analysis performed on the two different melts using a Bruker Q4 Tasman Advanced CCD-Based Optical Emission Spectrometer showed that both melts were in the range of standard A356 alloy chemistry values.Surface RoughnessThe results of the surface roughness testing are reported in Table 6. After confirming the data followed a normal distribution, a student’s t-Test with a level of significance (α) of 0.05 was used to determine if the surface roughnesses of the cylindrical specimens cast using different mold materials were significantly different.Table 6.

Surface roughness average (Ra) measurements of the overall metal specimensMold MaterialNo-Bake ViriCast™ ZCast®Mean (µm) 12.17 13.62 15.62SD (µm) 2.87 3.11 2.85T-Test Comparison p-ValueX 0.1223 0.0002----------- X 0.0559Specimens prepared using ZCast® molds had the roughest surface finish on average. The samples produced using no-bake molds were significantly smoother than those cast from ZCast® but not compared to ViriCast™. Additionally, the ViriCast™ 3DP and ZCast® molds produced significantly equivalent surface roughnesses.Surface finish is a function of sand particle size and distribution. Fine grain sands tend to produce better surface finishes but reduce the permeability of the mold to gasses [19]. Additionally, previous tests show that ZCast® molds produce a larger amount of gasses during casting due to the binder used during the binder jetting process [2].

Additionally, previous tests show that ZCast® molds produce a larger amount of gasses during casting due to the binder used during the binder jetting process [2]. The increase in gas in ZCast® molds, in combination with the smaller particle size in both 3D powders, could explain the larger surface roughness in ZCast® and ViriCast™, although significantly greater in ZCast® castings.Sand casting processes typically produce cast parts with surface roughness values between 12.5 and 25 µm [20]. Although specimens produced using the 3DP molds had a rougher surface finish than the no-bake specimens, their surface roughness values still fall on the low range of typical sand cast surface roughness values [20].DensityThe average densities of cylindrical specimens of A356-T6 aluminum cast from different mold materials are reported in Table 7. The densities of the specimens cast from 3DP molds were less than the standard density for the A356-T6 alloy (2.66-2.71 g/cm3) [21].

The densities of the specimens cast from 3DP molds were less than the standard density for the A356-T6 alloy (2.66-2.71 g/cm3) [21]. The sample densities were lower than expected due to porosity observed throughout the cast pieces. The overall density data for the specimens did not have a normal distribution; as a result a non- parametric Wilcoxon test (α = 0.05) was used for statistical comparison. The average density of no-bake and ViriCast™ castings didn’t vary significantly as well as between ViriCast™ and ZCast®. On the other hand, the densities of ZCast® castings did vary significantly from no-bake castings. This could be due to larger percentage of porosity in the ZCast® castings. The density did not vary significantly throughout the length of the specimens regardless of the mold material.Table 7.

This could be due to larger percentage of porosity in the ZCast® castings. The density did not vary significantly throughout the length of the specimens regardless of the mold material.Table 7. Average density measurements of overall metal specimens (mean values and SD)Mold MaterialNo-Bake ViriCast™ ZCast®Mean (g/cm3) 2.61 2.61 2.59SD (g/cm3) 0.05 0.02 0.04Wilcoxon Comparison p-ValueX 0.1497 0.0175------- X 0.1837PorosityThe amount of porosity present in the specimens was determined by analyzing micrographs of the polished aluminum samples. Micrographs demonstrating the porosity of metal cast using the three different mold materials are shown in Figure 4. The average porosity values for the entire samples are reported in Table 8. After determining the data was not normal, a non- parametric Wilcoxon test (α = 0.05) was used to determine if differences in the data were significantly different.Figure 4.

After determining the data was not normal, a non- parametric Wilcoxon test (α = 0.05) was used to determine if differences in the data were significantly different.Figure 4. Micrographs of T6-A356 aluminum cast in traditional no-bake (a), ViriCast™ (b), and ZCast® (c) moldsTable 8. Average porosity values of overall metal specimens (mean values and SD)Mold Material Mean (%) 0.65 1.13 1.59No-Bake ViriCast™ ZCast®SD Wilcoxon Comparison (%) 0.53 0.71 1.36p-ValueX <0.0001 <0.0001-------- X 0.1445The porosity observed in samples cast in ZCast® molds was higher than the samples cast in ViriCast™ molds, but not significantly. Porosity in the samples cast in ZCast® molds had a large standard deviation in relation to the average. The porosity seen in specimens prepared with no-bake molds was significantly less than the cylinders prepared using 3DP molds. This is most likely due to the higher binder content of 3DP molds.

The porosity seen in specimens prepared with no-bake molds was significantly less than the cylinders prepared using 3DP molds. This is most likely due to the higher binder content of 3DP molds. During the pouring process, off-gassing of the binder causes entrapped gasses that lead to porosity in the final cast parts.During the cylinder sectioning process, the orientations of the middle sections were not kept consistent. So, data from the middle sections may represent data taken from Faces 1 or 2 (see Figure 3b). The top data points were all taken at Face 1, and the bottom data points were either taken at Face 2 or Face 3. As a result, the data from the middle section does not provide useful information for comparison.

As a result, the data from the middle section does not provide useful information for comparison. The data from the top and bottom sections still can be used to analyze trends in metals properties throughout the length of the mold.Analysis of the top and bottom sections showed that porosity did not vary significantly throughout the length of the cylindrical samples cast in 3DP molds, as shown in Figure 5. The standard deviation was so large in both locations that even though there appears to be a trend in the means, the differences were not statistically significant. In the no-bake molds, the metal at the bottom of the mold was significantly more porous than the metal at the top of the mold.)%(y t i s o r o P4.50 4.00 3.50 3.00 2.50 2.00 1.50 1.00 0.50 0.00No-BakeVCastZCastTopBottomLocation in MoldFigure 5.

Variation of porosity of cast A356-T6 throughout the length of the moldMicrostructureAfter etching the polished aluminum samples, the dendrite arm spacing was measured using optical microscopy. Statistical analysis was performed on the data using a non-parametric Wilcoxon test (α = 0.05). The results of the microscopy measurements are shown in Table 9. Micrographs displaying representatives of the A356 microstructure from each mold type are shown in Figure 6. The microstructure was analyzed to determine if the different melts provided like mechanical properties for the aluminum cylinders. Finer dendrite arm spacing is desirable for better mechanical property performance. The larger the dendrite arm spacing, the coarser the microconstituents and the more prominent their effects on properties [22].Table 9.

The larger the dendrite arm spacing, the coarser the microconstituents and the more prominent their effects on properties [22].Table 9. Average dendrite arm spacing in the overall metal specimensMold Material Mean (µm) 41.2 72.0 72.9No-Bake ViriCast™ ZCast®SD Wilcoxon Comparison (µm) 5.32 9.74 7.28p-ValueX <0.0001 <0.0001-------- X 0.7559Figure 6. Dendritic microstructure of A356-T6 alloy cast in no-bake a), ViriCast™ b), and ZCast® c) moldsThe dendrite arm spacing in the samples cast in ZCast® and ViriCast™ molds were not significantly different. The samples prepared using no-bake molds had significantly smaller dendrite arm spacing than the 3DP prepared specimens. This indicates the heat treating processes varied between the two 3DP pieces and the no-bake parts.

This indicates the heat treating processes varied between the two 3DP pieces and the no-bake parts. The T6 heat treatment was performed on all the specimens cast in 3DP molds at the same time, but the heat treatment on the samples cast in no-bake molds was completed at a different time and in another furnace. Since the heat treatments differed, the mechanical properties (hardness and compressive yield strength) of the no-bake specimens cannot be compared to those of the 3DP prepared metal cylinders. Heat treating does not affect the porosity, surface roughness, or density of the aluminum specimens, thus valid comparisons can still be made between the no-bake cylinders and 3DP cylinders for the above tests [22].HardnessThe Vickers Hardness values of the cylindrical specimens were measured and used to compare the metals cast from different mold materials.

After determining the data was notnormal, a non-parametric Wilcoxon test (α = 0.05) was used to determine if differences in the data were significantly different. The results of the hardness testing are reported in Table 10.Table 10. Vickers hardness values of the overall metal specimensMold MaterialNo-Bake ViriCast™ ZCast®Mean (HV) 82.1 92.7 94.3SD Wilcoxon Comparison (HV) 4.83 9.67 9.60p-ValueX <0.0001 <0.0001---------- X 0.6204The Vickers hardness values for specimens produced using ViriCast™ and ZCast® molds did not vary significantly from each other. The test values for both specimens produced using 3DP molds fell within the normal hardness value range of 87.38 – 96.65 HV for the A356-T6 alloy [21]. The hardness values did not vary significantly throughout the length of the cast cylindrical specimens. The specimens produced by both the 3DP printed molds were significantly harder than the samples produced using traditional no-bake molds.

The specimens produced by both the 3DP printed molds were significantly harder than the samples produced using traditional no-bake molds. The differences observed in hardness between the specimens produced with no-bake and 3DP molds are most likely due to an issue with heat treating the no-bake specimens, as mentioned previously. In all of the mold types, hardness values did not vary significantly with mold location. The metal in the top and bottoms of the molds had statistically equivalent hardness values.Compression TestingMetal cylinders produced from the 3DP molds were machined to match compression specimen requirements [15]. Compressive yield strengths were determined and compared against published values to determine if the 3DP molds produced cast samples with mechanical properties comparable to traditional foundry techniques.

A Wilcoxon test (α = 0.05) was used to determine if there were any significant yield strength differences between castings from no-bake, ZCast®, and ViriCast™ molds. The results of the compression testing are shown in Table 11. An example stress-strain curve for a cylinder cast using a ZCast® mold is shown in Figure 7.Table 11. Compression testing of cast A356-T6 alloy species. No-bake compressive yield strengths were obtained from published sources [14].Mold Material No-Bake ViriCast™ ZCast®Mean Compressive Yield Strength, σ (MPa) 165-195 [22–24] 170.8 180.8SD Wilcoxon Comparison(MPa)p-Value------30.5 35.8X 0.2971350300) a P M( σ250200, s s e r t S150100ViriCast™ZCast®500024 Strain, ε (%)68Figure 7. Compressive stress-strain curve for A356-T6 cylinders cast in ZCast® and ViriCast™ moldsThe compressive yield strengths of the specimens cast using 3DP molds fell within the range of published data.

This shows that metal parts produced using additive manufacturing techniques have the same mechanical properties in compression as those produced using traditional sand casting techniques. The yield strengths of the cast metal did not vary significantly between the two 3DP mold materials (p=0.2971).4.CONCLUSIONS AND FUTURE WORKThe binder jetting 3DP process has been used to produce molds to cast complex structures. In this work, two 3DP powders, ViriCast™ and ZCast®, were compared on the basis of handleability of the printed molds and quality of the cast metal they produced. The two 3DP powders yielded nearly identical samples. Samples cast using ViriCast™ and ZCast® molds were statistically equivalent on all six tests performed.Additionally, the handleability and metal casting abilities of the 3DP powders were compared to traditional no-bake foundry sand molds.

It was determined that the binder jetting process can be used to produce metal specimens with similar properties to those created usingtraditional sand casting techniques. Both the ViriCast™ and ZCast® 3DP molds produced cast metal parts with the same mechanical performance and hardness as traditionally prepared A356- T6. However, ZCast® molds produced cast A356-T6 with greater surface roughness and decreased density than the samples prepared with no-bake molds. Furthermore, the no-bake molds had a significantly higher tensile strength than the 3DP molds, which made them more handleable, while produced castings with less porosity and smaller dendrite arm spacing.As technology advances, modeling, including flow modeling and solidification, will yield higher quality castings by minimizing porosity resulting in desired microstructure.

Due to the freedom of design provided by Additive Manufacturing, the molder has the ability to overcome the manufacturing constraints of traditional mold making in order to generate optimal complex castings. Continued work with these molding materials, in addition to others, will provide the ability to enhance existing lightweight, stiff cellular structures with designed mesostructure [6].REFERENCES[1]P. R. Beely, Foundry Technology. Butterworth-Heinemann, 2001.[2] D. A. Snelling, V. Tech, R. Kay, A. Druschitz, and C. B. Williams, “Mitigating GasDefects in Castings Produced from 3D Printed Molds,” in 117th Metalcasting Congress, 2012.[3]3D Systems, “Solutions: Metal Casting.” [Online]. Available: http://www.zcorp.com/en/Solutions/Castings-Patterns-Molds/spage.aspx.[4]ExOne, “ExOne: Digital Part Materialization.” [Online]. Available: http://www.exone.com/materialization/what-is-digital-part-materialization/sand.[5] Viridis3D, “Metal Casting.” [Online].

ficient concurrent toolpath planning for multi-material layered manufacturing S.H. Choi and W.K. Zhu Department of Industrial and Manufacturing Systems Engineering, The University of Hong Kong, Hong Kong Abstract Reviewed, accepted September 10, 2008This paper proposes an algorithm for planning efficient concurrent toolpaths to reduce the build-time of fabricating multi-material prototypes by layered manufacturing. The algorithm first sorts and partitions slice contours into hierarchical families of specific materials to enhance concurrent tool movements. It then detects overlapping of parametric tool envelopes with a “Two-phase Overlap Query Algorithm” to avoid potential tool collisions. Finally, it plans concurrent movements with an “Immediate Fabrication Algorithm” (IFA) to enhance fabrication efficiency by reducing idle time of tools. The algorithm is being implemented in a multi-material virtual prototyping system.

The algorithm is being implemented in a multi-material virtual prototyping system. It can be adapted for control of physical fabrication of multi-material prototypes when appropriate hardware becomes available.1 IntroductionToday’s market is characterized by intensifying global competition, increasing product complexity, continuous innovation dynamism, and harsh customer requirements. In order to survive and to maintain competitiveness, it is imperative for companies to strive to reduce the time-to-market and costs of product development.A significant amount of development time and costs are often spent on product design and process validation. For this purpose, prototypes are commonly used for verification and testing of aesthetics, processes, and functionalities. Fabrication of a prototype is however an expensive and time-consuming undertaking.Prototypes can be hard or soft.

Fabrication of a prototype is however an expensive and time-consuming undertaking.Prototypes can be hard or soft. A hard prototype is a model which can be physically touched and manipulated, while a soft one is digitally simulated and may be regarded as a computer graphical image of the product.A hard prototype is traditionally crafted by a skilled pattern-maker. It is slow and dimensionally inconsistent. About two decades ago, the stereolithography apparatus (SLA), generally known as the first commercial layered manufacturing (LM) or rapid prototyping (RP) machine, was developed. It is a relatively faster process that fabricates a physical prototype additively from a 3D CAD model layer by layer.

It is a relatively faster process that fabricates a physical prototype additively from a 3D CAD model layer by layer. LM systems are now widely used in product development and medical applications to save time and costs (Yildirim et al., 2006).On the other hand, virtual prototyping (VP) is a process of using a digital prototype, in lieu of a physical one, for testing and evaluation of specific characteristics of a product or a manufacturing process. Virtual prototypes can be sent via the Internet to customers to solicit comments, or the process parameters can be tuned for optimal fabrication of physical ones. ThusVP reduces the number of physical iterations and thereby the associated manufacturing overheads, leading to faster and cost-effective product development.Currently, most commercial LM machines can only produce single-material prototypes. However, there is an increasing need for multi-material prototypes (Malone and Lipson 2007).

However, there is an increasing need for multi-material prototypes (Malone and Lipson 2007). It would be very desirable and useful to develop multi-material layered manufacturing (MMLM) technology. The material-depositing mechanism may consist of an array of nozzles or tools, each of which deposits a type of material on specific areas of a layer. Such a mechanism could be adapted from available hardware, but its control would require a relatively intelligent software system to accomplish effective and efficient fabrication of multi-material prototypes. Therefore, the development of MMLM technology is largely a software issue.Indeed, processing of complex multi-material slice contours and planning of multi-toolpaths are particularly important issues of MMLM. Toolpath planning of MMLM may be sequential or concurrent. Concurrent toolpath planning saves the build-time of prototypes and hence shortens product development process.

Toolpath planning of MMLM may be sequential or concurrent. Concurrent toolpath planning saves the build-time of prototypes and hence shortens product development process. However, concurrent movements of tools may cause collision problems. It is desirable to develop an efficient concurrent toolpath planning algorithm for MMLM. This algorithm should allow as many concurrent tool movements as practicable without tool collision. It would be a significant step in development of MMLM for advanced product development.2 Related works and research gapToolpath planning involves mainly two issues, namely tool collision detection and tool motion planning. Collision detection, also known as interference detection or contact determination, is a multi-discipline issue which is particularly important in computer graphics, CAD/CAM, motion planning, etc. (Ericson, 2005). Accuracy of detection and speed of computation are always two contradicting factors.

(Ericson, 2005). Accuracy of detection and speed of computation are always two contradicting factors. The computational bottleneck of collision detection attracts many researchers. Many algorithms have been proposed and each of them has its merit and deficiency (Lin and Gottschalk, 1998).Many researchers, particularly in areas of robotics and manufacturing, have contributed a lot to collision detection/avoidance and motion planning. Fraile et al. (2006) developed three methods to facilitate on-line path planning for multi-manipulator systems in dynamic environments with collision avoidance, where multi-manipulators may move concurrently in their common workspace. Liu and Kulatunga (2007) developed a simultaneous task allocation and motion coordination approach, which can solve the scheduling, planning and collision avoidance problems simultaneously. Kok et al.

Kok et al. (2005) applied coordination graphs to a robotic domain by assigning tasks to the robots and then coordinating the different tasks. Ding et al. (2004) developed a global interference detection method for 5-axis machining of free-form surfaces. In order to simplify the process of updating tool positions and orientations in 5-axis machining, this method models a cutter and its holder by a hierarchy-oriented bounding box structure, whereasthe workpiece surfaces are approximated by an octree.In the domain of layered manufacturing, some research outputs of toolpath planning have been published. Zhu and Yu (2002) proposed a toolpath generation method of multi-material assembly for rapid manufacturing. With this method, a dexel-based “spatio-temporal” modelling approach was proposed for collision detection and concurrent tool movement. Lee and Kim (2006) proposed a multi-robot cooperation-based mobile printer system.

Lee and Kim (2006) proposed a multi-robot cooperation-based mobile printer system. A user uses this system to draw a picture on an input window of the GUI, and then the host computer commands client printer-robots to reproduce the same on a paper in a finite time. Malone and Lipson (2007) developed a set of software and hardware, called Fab@Home, which can sequentially fabricate multi-material prototypes of specific materials.However, most of these toolpath planning algorithms are either sequential or can only handle relatively simple multi-material prototypes. Choi and Cheung (2005) developed a multi-material virtual prototyping system and a topological hierarchy-based approach to toolpath planning for MMLM. They later enhanced the toolpath planning algorithm by a parametric tool-envelope method (2006). Based on sorting and grouping toolpath-sets of specific materials, this method can generate concurrent toolpaths for relatively more complex objects.

Based on sorting and grouping toolpath-sets of specific materials, this method can generate concurrent toolpaths for relatively more complex objects. However, there may be considerable idle time of tools and the algorithm assumes constant tool speed. This paper presents the work in attempt to further enhance this concurrent toolpath planning algorithm. It is aimed that the algorithm could schedule for as many concurrent nozzle movements as practicable, and could handle variable tool speed. The algorithm is expected to substantially shorten the build-time of MMLM and hence the product development process accordingly.

The algorithm is expected to substantially shorten the build-time of MMLM and hence the product development process accordingly. Moreover, the algorithm could be employed in other fields, like multi-robot motion planning and multi-arm process.robot assembly3 Demonstration of proposed algorithm 3.1 Previous worksThis section first demonstrates how the to topological hierarchy-sorting approach toolpath planning for MMLM works, and then presents the enhancements.Figure 1 is a CAD model of a discrete multi-material machine part. Using the slicing and sorting modules of our in-house system (Choi and Kwok, 2002), we first slice the part into some layers and sort the topological hierarchy relationship of slice contours. A slice of the model is shown in Figure 2. Contours are categorized into three levels (level 0 to level 2), from the outermost contours to the innermost ones.

A slice of the model is shown in Figure 2. Contours are categorized into three levels (level 0 to level 2), from the outermost contours to the innermost ones. For example, in envelope E7, which encloses oneFigure 1 A discrete multi-material machine partor more families of contours, Contour C9 (at level 0) is parent of C10 (at level 1) and they form a “contour family (pink)”.

C12 forms a contour family (yellow) by its own and so do C13, C14, and C15, and they are enclosed in E9, E10, E11, and E12 (all in yellow), respectively.C4E 3E 7C13E 8C11E10E12C5E 4C6E 5C9C10C15C7C8E 6E 9E11C12C14C3C1C2E 1E2Figure 2 A slice of the machine part in Figure 1 (Ei for envelope and Ci for contour)Topological hierarchy relationship of contoursParent-and-child list for contour containmentContour familiesToolpaths MaterialLevel 0C1C3 E2C6 C5 C4 E3 E4 E5C7C9C11E8Level 1C2 E1C8 E6C10Level 2C12 C13 C14 C15E7Group related toolpaths as toolpath-sets for association with a related nozzle based on material property: Set 1 (PC1,2+PC3) N1 (nozzle 1 for blue material) Set 2 (PC4+PC5+PC6 ) N2 (nozzle 2 for green material) Set 3 (PC7,8) N3 (nozzle 3 for red material) Set 4 (PC9,10+PC11) N4 (nozzle 4 for pink material) Set 5 (PC12+PC13+PC14+PC15 ) N5(nozzle 5 for yellow material)1.

C1 2. C3 3. C4 4. C5 5. C6 6. C7 C8 7. C9 C10 8. C11 9. C12 10. C13 11. C14 12. C15C2PC1,2 PC3 PC4 PC5 PC6 PC7,8 PC9,10 PC11 PC12 PC13 PC14 PC15Blue Blue Green Green Green Red Pink Pink Yellow Yellow Yellow YellowTable 1 The topological hierarchy and toolpath-sets of the contours in Figure 2A contour family of a specific material forms a toolpath. Inside E7, there are six toolpaths, namely PC9,10, PC11 (both in pink) and PC12, PC13, PC14, PC15 (all in yellow). Toolpaths for a specific material are grouped together to form a toolpath-set. For example, toolpaths PC9,10, and PC11 (both in pink) are grouped together to form toolpath-set 4 associated with nozzle N4. Table 1 shows complete grouping of toolpaths. The outermost contour of each contour family is offset by a distance of the tool radius to form tool envelope.

Table 1 shows complete grouping of toolpaths. The outermost contour of each contour family is offset by a distance of the tool radius to form tool envelope. A tool envelope, in which only a tool is allowed to move along the related toolpath-set to deposit the specific material, is constructed for detection of potential tool collisions. If a pair of tool envelopes overlaps, there may be tool collisions in the envelopes. To balance the accuracy and speed of overlap query, the shape of a tool envelope is parametric, in that it can be varied from being a quadrangle that results in the least accurate but the fastest overlap query, to being an exact offset of the contour that entails the most accurate but the slowest overlap query. Table 2 shows the result of overlap queries.

Based on this result, materials can be deposited concurrently by arranging the toolpath-sets into two groups, namely group 1 (blue, green, and pink) and group 2 (red and yellow), with the related nozzles associated accordingly as shown in Figure 3.Blue (Set 1) Green (Set 2) Red (Set 3) Pink (Set 4)Green (Set 2) Non-overlapRed (Set 3) Overlap Non-overlapPink (Set 4) Non-overlap Non-overlap Non-overlapYellow (Set 5) Non-overlap Non-overlap Non-overlap OverlapTable 2 Result of overlap queries of the previous algorithmStep 1: nozzles N1, N2, and N4 deposit the specified materials concurrently and then return to the datum point.Step 2: nozzles N3, and N5 deposit the specified materials concurrently and then return to the datum point.N1 (blue) for Set 1 (PC1,2+PC3) N2 (green) for Set 2 (PC4+PC5+PC6 ) N4 (pink) for Set 4 (PC9,10+PC11)N3 (red) for Set 3 (PC7,8) N5 (yellow) for Set 5 (PC12+PC13+PC14+PC15 )Figure 3 Concurrent nozzle movements by the previous algorithmN2N3N 1N4N5t<=8, nozzles N1, N2, N4 deposit blue, green, and pink materials concurrently8<t<=12, nozzles N3, N5 deposit red, and yellow materials concurrentlyFigure 4 Virtual fabrication of the machine part by the previous algorithmFigure 4 shows the virtual fabrication of the machine part by the previous algorithm.

Unless specified otherwise, “t” stands for time and is in “time unit”. Figure 5 shows the time chart of the previous algorithm. It can be seen that there is some idle time of tools. Take nozzle N3 (red) and N1 (blue) for example. In Figure 2, E6 (red) overlaps with E1 (blue). Since E1 is grouped into toolpath-set 1 with E2, nozzle N3 (red) cannot start to deposit material until N1 (blue) finishes depositing both E1 and E2. In other words, the time spent on depositing E2 (blue) is the idle time of N3 (red). In regard to nozzle N5 (yellow), depositing finishes immediately depositing material in E7 (pink).

In other words, the time spent on depositing E2 (blue) is the idle time of N3 (red). In regard to nozzle N5 (yellow), depositing finishes immediately depositing material in E7 (pink). However, N5 is grouped into toolpath-set 5 which is triggered in step 2, so it has to wait until the finish of toolpath-set 1 (step 1), that is, until N1 (blue) finishes depositing E2 (blue).itshould after N4start(pink)Figure 5 Time chart of previous algorithmIn order to reduce idle time of tools, we propose an enhanced concurrent toolpath planning algorithm called “Immediate Fabrication Algorithm” (IFA). This algorithm does not group toolpaths into toolpath-sets. Instead, it increases the number of tool envelopes (hence smaller envelopes) such that when a contour family (toolpath) in an envelope is finished, deposition of another contour family (toolpath) in the corresponding overlapping envelope can be more likely started immediately.

However, due to increased number of envelope overlap queries, it is necessary to develop an efficient overlap query algorithm first, which is introduced in the following section.3.2 Two-phase Overlap Query Algorithm (TOQA)There are twelve envelopes in Figure 2. Pairwise overlap query of n envelopes requires (n-1)+(n-2)+……+2+1=n(n-1)/2=O[n(n-1)/2] tests in the worst case. Due to the quadratic time complexity, full enumeration of testing all envelope pairs becomes too expensive even for a moderate value of n. To speed up the process, the number of pairwise tests must be reduced. When a slice of a prototype is complex and the number of tool envelopes is big, time complexity of overlap query is lengthy, impairing the performance of real-time virtual prototyping simulation.

When a slice of a prototype is complex and the number of tool envelopes is big, time complexity of overlap query is lengthy, impairing the performance of real-time virtual prototyping simulation. To tackle this problem, a “Two-phase Overlap Query Algorithm” (TOQA), which consists of “material-group phase” and “envelope phase”, is proposed as follows:for layer Li { “material-group phase” overlap query; if (overlap detected) “envelope-phase” overlap query; }// The “Two-phase Overlap Query Algorithm”When executing overlap query for a layer, the “material-group phase” is invoked first. It groups envelopes of the same material property together and then executes pairwise overlap query between material-groups. Time complexity for m materials is O[m(m-1)/2], where m is usually much smaller than n.

Time complexity for m materials is O[m(m-1)/2], where m is usually much smaller than n. If overlapping is detected during the “material-group phase”, the “envelope phase” will be invoked to execute further pairwise tests between envelopes involved.Now let’s take Figure 2 for example and analyze the improved efficiency of the proposed algorithm. In Figure 2, there are n=12 tool envelopes for m=5 materials. When executing overlap queries of the layer, the “material-group phase” is invoked first. Pairwise tests between material-groups are executed and the time complexity is O[m(m-1)/2]=O[5*(5-1)/2]=O(10).

Pairwise tests between material-groups are executed and the time complexity is O[m(m-1)/2]=O[5*(5-1)/2]=O(10). The result of overlap queries of the material-group phase is shown in Table 3.Blue Green Red PinkGreen Non-overlapRed Overlap Non-overlapPink Non-overlap Non-overlap Non-overlapYellow Non-overlap Non-overlap Non-overlap Overlapmaterial-group phaseenvelope phaseE1 (Blue) E6 (Red) OverlapE2 (Blue) Non-overlapE9 (Yellow) E10 (Yellow) E11 (Yellow) E12 (Yellow)E7 (Pink) Overlap Overlap Overlap OverlapE8 (Pink) Non-overlap Non-overlap Non-overlap Non-overlapTable 3 Result of TOQA for Figure 2The “material-group phase” overlap query algorithm finds that there are overlaps between material-groups blue and red, and between material-groups pink and yellow. So, further pairwise tests between envelopes involved are executed in the “envelope phase”.

So, further pairwise tests between envelopes involved are executed in the “envelope phase”. For material-groups blue and red, the time complexity of pairwise tests for three envelopes is O[3*(3-1)/2]=O(3) while the test between E1 and E2 (of the same material) can be ignored. So here, the time complexity is reduced to O(3-1)=O(2) and overlapping of E1 (blue) and E6 (red) is found out as shown in Table 3 (envelope phase, left). For material-groups pink and yellow, there are two pink envelopes and four yellow envelopes. All-pairwise overlap query requires O[6*(6-1)/2]=O(15) tests. Ignoring a test between two pink envelopes (of the same material) and O[4*(4-1)/2]=O(6) tests between four yellow envelopes (of the same material), the time complexity is reduced to O(15-1-6)=O(8). It is found that E7 (pink) overlaps with E9, E10, E11, E12 (all are yellow), as shown in Table 3 (envelope phase, right).

It is found that E7 (pink) overlaps with E9, E10, E11, E12 (all are yellow), as shown in Table 3 (envelope phase, right). Together, the time complexity of the “envelope phase” for Figure 2 is O(2+8)=O(10).Totally, the time complexity of the two-phase test is O(10+10)=O(20). Compared with that of full enumeration of testing all envelope pairs, O[n(n-1)/2]=O[12*(12-1)/2]=O(66), we achieve an improved efficiency of (66-20)/66=69.7%, as shown in Figure 6. Generally, since m is smaller than n, O[m(m-1)/2] is much smaller than O[n(n-1)/2]. Though further tests are required in “envelope phase”, the complexity is significantly reduced. Meanwhile, ignoring the same tests between envelopes of materials can further reduce the time complexity. Indeed, the more contour families in a slice, the better efficiency of the proposed “Two-phase Overlap Query Algorithm”.

Indeed, the more contour families in a slice, the better efficiency of the proposed “Two-phase Overlap Query Algorithm”. This is particularly important for virtual prototyping simulation.Figure 6 Comparison of complexities of full test and TOQA3.3 Immediate Fabrication Algorithm (IFA)To reduce idle time of tools, we propose the “Immediate Fabrication Algorithm” (IFA). This algorithm does not group toolpaths into toolpath-sets. Instead, it improves concurrency of tool movements by enhancing the likelihood of triggering material deposition by the subsequent tool immediately after one finishes the contours in the overlapping envelope.

The psuedocode of IFA is as follows:IFA ( ) // Immediate Fabrication Algorithm (IFA) { for layer Li { for envelope Ei { Two-phase Overlap Query; } group non-collision toolpaths; while (layer not finished) { display concurrent toolpaths; check flag of envelope status; if (overlapped envelope finished) { add a new nozzle to current toolpaths; trigger the new nozzle; } } //end of “while (layer not finished)” } // end of “for layer Li” } // end of “IFA ( )”Take nozzle N3 (red) and N1 (blue) for example again. In TOQA, it is found that envelope E6 (red) overlaps with E1 (blue). So nozzle N3 (red) stands by. When t=3 (Figure 7), N1 (blue) has just finished depositing in E1 (blue) and continues to E2 (blue). Since E6 (red) does not overlap with E2 (blue), right at this moment (t=3), nozzle N3 (red) is triggered to deposit material ontimecontour family in envelope E6 (red) immediately.

Since E6 (red) does not overlap with E2 (blue), right at this moment (t=3), nozzle N3 (red) is triggered to deposit material ontimecontour family in envelope E6 (red) immediately. In regard to nozzle N5 (yellow), envelopes E9, E10, E11, and E12 (all in yellow) overlap with E7 (pink) in TOQA, so nozzle N5 stands by. When t=4 (Figure 8), N4 (pink) has just finished depositing material in envelope E7 (pink) and continues to deposit material on the contour family in E8 (pink). Since E9, E10, E11, and E12 (all in yellow) do not overlap with E8 (pink), right at this moment (t=4), nozzle N5 (yellow) is triggered to deposit material on contour families in envelopes E9, E10, E11 and E12 (all in yellow) immediately. N3N2N3N2N4N4N 1N 1N5Figure 7t=3, N1, N2, N4 moveFigure 8t=4, N1, N2, N3, N4 moveFigure 9 t=8, layer completedconcurrently. N1 has just finished E1concurrently.

N1 has just finished E1concurrently. N4 has just finished E7 (pink)(blue) and continues to E2 (blue); N3and continues to E8 (pink); N5 is triggeredis triggered to start E6 (red)to start E9, E10, E11 and E12 (all in yellow)Figures 7 to 9 show the complete fabrication process of IFA, while Figure 10 is the corresponding time chart. Comparing Figure 10 with Figure 5, it can be seen the build-time of the previous algorithm is twelve units while IFA gives eight, improving the efficiency by 33%, as shown in Figure 11.Figure 10 Time chart of IFA Figure 11 Comparison of build-times of theprevious algorithm and IFA4 Case studyFigure 12 shows a CAD model of a gearbox assembly and a slice of it. There are eight contours and four types of materials. The topological hierarchy relationship and the result of TOQA are shown in Tables 4 and 5, respectively.

There are eight contours and four types of materials. The topological hierarchy relationship and the result of TOQA are shown in Tables 4 and 5, respectively. Comparing full enumeration test with the TOQA, time complexity of TOQA is only O(8) while full test takes O(21) as in Figure 13, improving the efficiency by 61.9%, which would be beneficial for real-time virtual prototyping simulation. Based on the result of TOQA, concurrent toolpaths are planned according to IFA.First (t=0), nozzle N1 (blue), N2 (purple), and N3 (brown) start to deposit the related materials in the specified envelopes concurrently (Figure 14). At t=3 (Figure 15), N1 (blue) has not finished task while N3 (brown) has just finished envelope E5 (brown) and continues to E6 (brown). At this moment, since E7 (green) does not overlap with E6, N4 (green) is immediately triggered to start depositing material in E7 without having to waiting for N1 to finish.

At this moment, since E7 (green) does not overlap with E6, N4 (green) is immediately triggered to start depositing material in E7 without having to waiting for N1 to finish. Figures 14 to 16 show the IFA process.Figure 12 A CAD model a gearbox assembly and slice of itTable 4 Topological hierarchy relationship of contours in Figure 12PurpleBrownGreenBlue Non-overlap Non-overlap Non-overlap Non-overlap Non-overlap Purple Overlapverlap BrownE5 ( Brown) E6 (Brown)E7 (Green) Overlap Non-overlapmaterial-group phase envelope phase Table 5 Result of Two-phase Overlap Query Algorithm for Figure 12Figure 17 shows the time charts of the previous algorithm and IFA. N4 (green) is triggered to deposit material in envelope E7 (green) as soon as E5 (brown) is finished, without having to wait for N1 (blue) to finish. Here, the build-time of IFA is five units.

N4 (green) is triggered to deposit material in envelope E7 (green) as soon as E5 (brown) is finished, without having to wait for N1 (blue) to finish. Here, the build-time of IFA is five units. In comparison with seven units of the previous algorithm, there is an improved efficiency of 28.6%, as shown in Figure 18.Figure 13 Comparison of time complexities of full test and “TOQA”N2N4N3N3N1N1Figure 14 t=0, N1, N2, and N3 start moving concurrentlyFigure 15 t=3, N1 has not finished task; N3 has just finished envelope E5 (brown) and continues to E6 (brown); N4 is triggered to start E7 (green)Figure 16 t=5, layer completedFigure 18 Comparison of the build-times of the previous algorithm and IFAPrevious algorithm5 Conclusions and future worksThis paper proposes an algorithm for improving the efficiency of concurrent toolpaths to reduce the build-time of MMLM prototypes.

Based on the topological hierarchy of slice contours, it adopts the “Two-phase Overlap Query Algorithm” to speed up overlap query of parametric tool envelopes for detection of tool collisions. Subsequently, the “Immediate Fabrication Algorithm”, is used to generateIFA Figure 17 Time charts of the previous algorithm and IFAefficient toolpaths to enhance concurrent movements by reducing idle time of tools. The algorithm is being implemented in a multi-material virtual prototyping system. It can be adapted for control of physical fabrication of multi-material prototypes when appropriate hardware becomes available. Currently, a tool envelope is constructed by offsetting the slice contour. If a pair of tool envelopes overlaps, tools in the tested envelopes may collide. This method is relatively a rough approximation. Future work would be dedicated to constructing envelopes around the tools, rather than around the slice contours.

This method is relatively a rough approximation. Future work would be dedicated to constructing envelopes around the tools, rather than around the slice contours. Collision detection between tool envelopes is executed at a certain time-interval while the tools are moving around. Direct motion planning of multi-tools would provide more accurate collision detection and improve the concurrency of tool movements. Moreover, in practical fabrication, it may be necessary to adjust the deposition speed for materials of different properties. These issues are generally concerned with “spatio-temporal” collaboration problems, which will be investigated in the near future.6 References Choi, S.H. and Cheung, H.H. 2005. A multi-material virtual prototyping system. Computer-Aided Design, 37(1):123-136. Choi, S.H. and Cheung, H.H. 2006. A topological hierarchy-based approach to toolpath planning for multi-material layered manufacturing. Computer-Aided Design, 38(2):143-156. Choi, S.H.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEfficient fabrication of Ti6Al4V alloy by means of multi-laser beam selective laser meltingFangzhi Li*, Zemin Wang* and Xiaoyan Zeng*Huazhong University of Science and Technology, Wuhan 430074, PR ChinaAbstractA self-developed four-laser beam selective laser melting (SLM) system was used to fabricate Ti6Al4V alloy samples in this study. The relative density, micro-hardness and mechanical properties of all isolated processing areas were compared under optimized processing parameters to ensure the consistency of this system. Microstructures in overlap areas are dominated by columnar grains along the building direction and matensitic needles αˊ inclined at about ± 45° to the building direction, which are similar with those in isolated areas. Mechanical properties in overlap areas are also not inferior to those in isolated areas.

Mechanical properties in overlap areas are also not inferior to those in isolated areas. The results prove the feasibility to fabricate large-scale components with a uniform microstructure and mechanical property by this SLM system. By the use of four lasers, this system can provide a high building rate of 80 cm3/h.KeywordsMulti-laser beam; Selective laser melting; Ti6Al4V; Microstructure; Mechanical property;IntroductionLaser additional manufacturing (LAM) has roused extensive attention for its capability to build up complex three dimensional features in an additive way. This process uses 3D CAD data as a digital information source to create three-dimensional metal components through fusing fine metal powders by a focused high-power laser beam layer by layer. Through this process, direct energy deposition or powder bed fusion are commonly used for manufacturing metallic materials.

Through this process, direct energy deposition or powder bed fusion are commonly used for manufacturing metallic materials. In the past few decades, researchers mainly focused on relative density, microstructures, surface properties and mechanical properties, etc. of different kinds of materials, size and shapes of metal components. Benefit by their research, LAM process was widely available in the field of aerospace, die forming and medical implants, etc., being as an effective complement of traditional processing methods like casting, forging and welding.However, the state-of-the-art process is not fully competent to series production for the demand of large-scale components and the restriction of fabrication efficiency.

In order to improve this efficiency, it is indispensable to increase the build rate significantly by matching a thicker powder layer with a higher scanning velocity under higher laser power[1].Laser cladding deposition (LCD), also referred to as direct laser deposition (DLD), laser melting deposition (LMD) or Laser-Engineered Net Shaping (LENS), which belongs to the direct energy deposition category, showed a great potential to form large scale components efficiently for its natural high laser power and mm-scale laser beam diameter. During this process, metal powder is injected into the melting pool to be melt and then solidified in a sudden97time onto a substrate. Tian[2] et al., achieved a powder delivery rate of 100 cm3/h Ti-4Al-1.5Mn titanium alloy by the use of a 5 kW CO2 laser with a laser beam diameter of 5 mm. Zhu[3] et al., further increase the powder delivery rate to 600-1000 g/h of Ti-6.5Al-3.5Mo-1.5Zr-0.3Si titanium alloy by a 8 kW CO2 laser.

Zhu[3] et al., further increase the powder delivery rate to 600-1000 g/h of Ti-6.5Al-3.5Mo-1.5Zr-0.3Si titanium alloy by a 8 kW CO2 laser. Yu[4] et al., obtained Ti6Al4V part with relative density up to 99.8% by the use of a 7 kW IPG fiber laser with spot size of 1.2 mm. Qiu[5] et al., successfully fabricated large Ti6Al4V structures by the use of a 4 kW disc laser and an automatic spot change collimator (from 0.2 to 6 mm). Their research demonstrate the capability of LCD process to fabricate large scale titanium structures efficiently. Due to these characteristics, LCD process has been presented a rapid expansion in aerospace field. On the other hand, large laser spot size and powder layer thickness will inevitably lead to poor shape precision together with the factor of direct energy deposition category.

On the other hand, large laser spot size and powder layer thickness will inevitably lead to poor shape precision together with the factor of direct energy deposition category. Therefore, LCD is still hard to meet the requirements of components with high precision and complex structure.Compared with LCD, selective laser melting (SLM), which belongs to the powder bed fusion category, is more suitable for precision forming of complex components with the aid of support structures and pre-paved powders. However, current SLM technology usually contains laser power no more than 500 W and powder thickness of 20-60 μm[6][7]. The related low building rate obviously cannot meet the demand of large scale components. In order to improve this efficiency, a kW-level laser was first integrated into a SLM system by Schleifenbaum[8] et al., together with a new multi-beam to maintain accuracy and detail resolution of manufactured components.

With this SLM system, Buchbinder[9] et al., increase the build rate to approx. 75.6 cm3/h for the production of AlSi10Mg parts while reaching densities above 99.5%. Ma[1] et al., further increase laser power to 6kW which can supply a sufficient energy density to obtain near-full-density part while maximum layer thickness reach to 150 μm of 1Cr18Ni9Ti stainless steel. The development of high power (HP) SLM provides a greater build rate than traditional SLM and a higher fabrication accuracy than LCD. However, limited by machining dimension, HP SLM is not capable to fabricate components in large scale size (usually limited in 250 mm × 250 mm).LCD and HP SLM both based on high power laser and its related large beam diameter and higher layer thickness, which in turn lead to restriction of poor accuracy in a certain extent.

In order to combine the advantage of high precision of traditional SLM process and high efficiency of high power LAM process, a multi-laser beam SLM system is developed. This system contains a multi-beam system including 4 independent continuous wave IPG YLR-500 fiber lasers together with 4 optical systems. In this multi-beam system, four lasers are in action simultaneously, providing four times the build rate compared with single laser SLM system while obtaining considerable accuracy. Meanwhile, this system can achieve a build envelope of 500 × 500 × 530 mm³. In the previous attempt, components has a great deal of cracking tendency at slicing line when overlap width is 0 for the lack of fusion. In order to overcome this shortage, overlap width from 0.5 to 2.5 mm was introduced in the experiments.

In order to overcome this shortage, overlap width from 0.5 to 2.5 mm was introduced in the experiments. In this situation, overlap area will go through another thermal cycle due to the re-melting of adjacent multi- beams.There is no doubt that the multi-beam SLM system can provide higher processing efficiency and the ability to process larger scale of components. However, the experience of fabrication by multi-beam SLM system is not been reported yet. The formability of this kind of system still remains to be verified.98In this paper, detailed investigation on the effects of overlap area between adjacent processing areas of Ti6Al4V titanium alloy by multi-beam SLM system was studied. The microstructure and mechanical properties of the fabricated samples in both isolated region and overlap area were compared.Self-developed multi-laser beam SLM systemA self-developed multi-beam SLM system (NRD-SLM-500) was used to process Ti6Al4V titanium alloy.

This system mainly consisted of four continuous wave IPG YLR-500 fiber lasers (λ=1.07 μm, maximum output power of 500 W, 100 μm spot size), four matrix-arranged scanning galvanometers (scanned area of 250 mm in both length and width), a building platform and a powder providing system, shown in Fig.1(a). By combining these matrix-arranged scanning galvanometers, four lasers can work together to build one big part or work separately to build several parts within 500(L)×500(W)×530(H) mm³. Compared with SLM systems with single laser beam, NRD-SLM-500 has two functions: larger build envelope size and higher build rate.The build rate can be improved by three methods: the use of four lasers, double-directionpowder delivery method and segmented powder delivery strategy, shown in Fig.1 (b).Fig.

1 Schematic diagram of NRD-SLM-500 system used in this studyThe manufacturing process can be decomposed as following steps shown in Fig.2: part design-file import-slicing and segmentation into four parts-fabricating. By the use of four lasers, this system can provide a high build rate of 80 cm3/h (considering the time of powder delivery).Experimental detailsGas atomized Ti6Al4V powders with an average particle size below 50 μm were used as the starting material, whose morphology is shown in Fig.3. Meanwhile, the chemical composition and particle size distribution were listed in Table 1. Based on the standard alternating x/y scanning strategy[10], relative hatch angle of scanning directions in adjacent scanning areas should be took into account. The sampling position, scanning strategy and configuration of tensile samples built in isolate and overlap area were shown in Fig.4 (a)-(c), respectively.

The sampling position, scanning strategy and configuration of tensile samples built in isolate and overlap area were shown in Fig.4 (a)-(c), respectively. In order to overcome the cracking tendency at slicing ling, overlap width from 0.5 to 2.5 mm was invested in the experiments.99Fig. 2 Manufacturing process of multi-laser beam SLM systemFig. 3 Characteristic morphology of Ti6Al4V powders used in SLM experiments Table 1 The chemical composition and particle size distribution of Ti6Al4V powders used in SLM processChemical composition (Wt%) Particle size distribution (μm)Ti 89.77Al 6.97V 3.26Dv10 19.0Dv50 33.8Dv90 57.3In the initial stage of the experiment, cuboid and tensile samples were built in 4 isolated processing areas to test the consistency. For each processing area corresponds the same types of laser and optical system, sample fabricated in 4 areas can be considered under the same processing conditions.

For each processing area corresponds the same types of laser and optical system, sample fabricated in 4 areas can be considered under the same processing conditions. In this paper, samples fabricated in area 3, area 4 and their overlap area were chosen as the research object. The SLM parameters used in this study are presented in Table 2.Surface morphology and microstructure were observed by a digital microscope (Keyence VHX-1000) and a scanning electron microscope (FEI Nova NanoSEM 450), Vickers hardness100tests were carried out using a HVS-1000 microhardness tester at a load of 9.8 N and a dwelling time of 20 s. Tensile test pieces were designed according to Chinese GB/T 228-2002 standard, as shown in Fig. 3 (a). For a further instruction, tensile test pieces fabricated at overlap area were pre-machined groove structure to test mechanical properties of overlap area.

3 (a). For a further instruction, tensile test pieces fabricated at overlap area were pre-machined groove structure to test mechanical properties of overlap area. All tensile test pieces were tested using a zwick/roell tester to evaluate the tensile properties at room temperature with a rate of 2 mm/min.Fig. 4 (a) sampling position and (b) laser scanning strategy, and (c), (d) configuration of tensile samples built in isolate and overlap areaTable 2 SLM manufacturing parameters used in this studyProcessing parameters Laser power (W) Scanning speed (mm/s) Scanning distance (mm) Layer thickness (μm) Hatch angle (°) Relative hatch angle (°) Overlap width (mm)Values 300 1000 0.12 40 90 0, 90 0.5-2.5Results and discussions3.1 Consistency of isolated processing areasTable 3 lists the relative density, micro-hardness and tensile properties of samples built in 1#-4# processing areas.

Under optimized parameter, unique micro-hardness and mechanical properties were obtained at a high building rate of 80 cm3/h101Table 3 Relative density, micro-hardness and tensile properties of samples built in 1#-4# processing areasRelative density (%) Micro-hardness (HV) Rm (MPa) Rp0.2 (MPa) El (%)1# area 99.92 408±17 1226±38 1131±41 5.3±0.52# area 99.81 407±16 1224±15 1140±11 5.2±0.63# area 99.84 397±11 1224±11 1151±39 5.3±0.64# area 99.93 411±4 1234±32 1127±24 5.03.2 Surface morphology of overlap areaFig. 5 shows the top surfaces of the samples that were fabricated across overlap area with varied overlap width and relative hatch angle. Fig.5 (a)-(e) represent relative hatch angle of 0 while Fig.5 (f)-(j) represent relative hatch angle of 90°. The white arrows show the scanning direction of final layer. The yellow ones show stair-steps produced in multi-beam SLM process. Apparently, stair-step effect is more serious when relative hatch angle is 0° than 90°.

The yellow ones show stair-steps produced in multi-beam SLM process. Apparently, stair-step effect is more serious when relative hatch angle is 0° than 90°. Normally, the stair-step effect will cause nonuniformity of actual powder layer thickness and do harm to SLM process. Fig.5 (k) and (l) show the three-dimensional contours of Fig.5 (b) and (g) for a more intuitive observation. The height of stair-step is only affected by the relative hatch angle but has no effect on overlap distance varied from 0.5 mm to 2.5 mm through the measurement of cross section from digital microscope. Specifically, the height difference decrease from approximate 180 μm to 60 μm by adjusting the relative hatch angle from 0 to 90°.Comprehensive consideration processing efficiency and convenience of study, overlapwidth of 1 mm and relative hatch angle of 90° were adopted in the following studies.Fig.

5 Surface morphology of overlap area1023.3 Microstructure and mechanical performance of multi-beamFig.6 shows the microstructure of samples built in 3#area, 4#area and overlap area. It can be seen from Fig.6 (a)-(c) that all built samples are dominated by columnar grains which tend to elonate along the building direction (i.e. Z direction). All these columnar grains grow over several layers and reach to mm-scale in length with the width of more or less 200 μm. Within the columnar grains, matensitic needles αˊ are observed for the fast melting-solidification process. Due to the fast cooling, the β phases transform to αˊ phase according to Burgers relation given by Equatation(1)[11]. These matensitic needles show an obvious herringbone distribution and inclined at about ± 45° to the building direction, indicated by blue slashes.

These matensitic needles show an obvious herringbone distribution and inclined at about ± 45° to the building direction, indicated by blue slashes. This incline is because of a specific Burgers relation between α/αˊ phase and β phase which dictates the αˊ growth orientation during fast melting-solidification process in SLM fabrication.(1)(1 1 0)𝛽𝛽 ↔ (0 0 0 1)𝛼𝛼< 1 1� 1 >𝛽𝛽 ↔ < 1 1 2� 0 >𝛼𝛼 Fig.(d)-(f) show the horizental section of samples built in 3#area, 4#area and overlap area, respectively. Chessborad patterns of spacing about 120 μm indicated by white dotted box can be observed for the reason of 90° hatch angle betweem adjacent layers in both 3# area, 4# area and overlap area.

Chessborad patterns of spacing about 120 μm indicated by white dotted box can be observed for the reason of 90° hatch angle betweem adjacent layers in both 3# area, 4# area and overlap area. The spacing width corresponding to the scanning distance used in the experiment.Comprehensive consideration of both vertical and horizental sections, metallographic structure and distribution in overlap area have no significant difference of the superposition of heat resource by the introduced laser beam compared with 3# and 4# area.Fig. 6 Microstructure of (a)-(c) vertical section and (d)-(f) horizontal section of samples built in 3#area, 4#area and overlap area, respectively.Fig.7 (a) shows results for Ti-6Al-4V powder, as-fabricated samples built in 3# area and overlap area. It can be seen that both powders and built samples show the hcp-Ti characteristics103of crystallography.

It can be seen that both powders and built samples show the hcp-Ti characteristics103of crystallography. This suggests that neither process of fabrication in isolated area nor re- melting in overlap area contains α/αˊ→β transformation. However, a slight shift in the hcp-Ti phase peaks to larger 2θ angles was observed, which may be by the reason of the solute redistribution, shown in Fig.7 (b).Fig. 7 (a) XRD result for Ti6Al4V powders, as-fabricated samples built in 3# area and overlap area, the green dotted box shows detailed information of 2θ from 34°-42°, (b) micro-hardness of overlap area and its adjacent processing areas and (c) mechanical properties in overlap area compared with isolated areas.3.4 Fabrication of large scale of component by multi-laser beam SLMFig. 8 large scale of component fabricated by multi-laser beam SLM system.

8 large scale of component fabricated by multi-laser beam SLM system. (a) fabrication process, (b) real figure of built component and (c) detailed drawing of (b).Based on the analysis of the above, 4 isolated processing areas and their related overlap areas have the uniform processing ability. This provides the premise to fabricate large scale of component with uniform performances. Fig.8 (a) shows the production process of one Ti6Al4V104model and Fig.13 (b) and (c) show the real figure of this model. Profit from 4 laser beams work at the same time, the multi-laser beam SLM system has an incomparable processing efficiency compared with traditional SLM systems. The largest diameter of this component reach to 401 mm and the height is 93mm.ConclusionIn this work, cuboid and tensile samples of Ti6Al4V were built in both isolated processing areas and overlap areas.

Through the experimental results, samples in 4 isolated areas and overlap areas have uniform microstructure and mechanical properties. Microstructures and mechanical properties of overlap areas are not affected by the relative hatch angle and overlap width. Considering the surface accuracy, θr is adopted in multi-laser beam SLM process. Under the work of 4 laser beams together, a large-scale of component can be fabricated in an efficiency way (build rate up to 80 cm3/h) finally.AcknowledgementsThis work is supported by the National Program on Key Basic Research Project of China (973 Program) under grant (no. 613281), the National Natural Science Foundation of China through program (no. 50905068), and the Fundamental Research Funds for the Central Universities through program (no. HUST: 2016YXZD005).Reference [1] Mingming Ma, Zemin Wang, Ming Gao, Xiaoyan Zeng. Layer thickness dependence of performance in high-power selectivelaser melting of 1Cr18Ni9Ti stainless steel.

However, both the CAD software and AM machine accept the input of more convenient and high-precision mathematical curve expressions. A predefined curved scanning pattern can be achieved in a galvanometer scanning system of selective laser melting (SLM) process, which can improve the scanning accuracy and efficiency for the shapes that can be represented by mathematical functions, e.g., circle. Therefore, this study proposes a layer construction-based method to directly generate the toolpaths for cellular structure with large file size. To demonstrate the proposed method, a case study on the toolpath generation for a flame arrestor element with a large quantity of curved fine channels, is presented.

To demonstrate the proposed method, a case study on the toolpath generation for a flame arrestor element with a large quantity of curved fine channels, is presented. Compared with the conventional printing preparation methods, the proposed method reduces up to 90% of the total preparation time and achieves higher quality toolpaths.IntroductionBenefiting from a layer-by-layer manufacturing way, AM process can deal with the problem of the fabrication of freeform geometry [1]. Nowadays, more and more multi-functional, and high-performance structures with extremely complex geometries are designed and manufactured via AM process [2~4]. However, when the issue comes to a massive fine porous structure, the shortcomings of the current AM process are showing. Firstly, limited by stereolithography (STL) file format, it is difficult for current data processing in AM chain to balance the accuracy of printing toolpaths and the efficiency of data transformation [5].

In general, the design software and the AM pre-processing software cannot handle such a large amount of data. However, the model itself can be described by some simple mathematical equations or other representations [6]. Secondly, as the design stage separating from the printing stage, it is hard to ensure whether the designed model can be manufactured with success or not at the first time [7]. Although design for AM (DfAM) is proposed to integrate AM constraints in the design processing [8~9], designers translate the AM processing information to three-dimensional (3D) geometric constrains [10~12] , which has a big difference with the printing toolpaths. Due to this difference, it is inaccurate to evaluate the manufacturability of the designed model for AM processing only1 Corresponding author. E-mail address: yicha.zhang@utbm.fr1202through its geometry. Both the printing toolpaths and AM parameters are needed to analyze.

E-mail address: yicha.zhang@utbm.fr1202through its geometry. Both the printing toolpaths and AM parameters are needed to analyze. In addition, the ability of the AM process is limited by the design and the AM pre- processing software. The generating algorithm of the toolpaths used in AM pre-processing has its limitations in dealing with small corner [13] and thin-wall region, especially in the porous structures.Therefore, to meet the future demands for multi-functional and high-performance structures, especially structures with large quantity of fine pores, this paper proposed an efficient layer construction method to generate accurate printing toolpaths for the selective laser melting process, which integrates the design stage and the printing stage.

The remainder of this paper is organized in the following: Section 2 presents the proposed method, Section 3 details the method with a case study of the design and fabrication of a flame arrester element, and Section 4 concludes this paper with perspectives on the future work.Method2.1. Method overviewFigure 1. Framework of layer construction method integrated design and manufacturing.The traditional data processing chain of AM includes model design, STL file creation, and build preparation[9]. It needs to transform the original CAD model to STL model, and the STL model to slicing model. For periodic cellular structure, this processing method usually causes the loss of printing accuracy and consumes a lot of time. Aiming to balance the accuracy and efficiency, the proposed method integrated design stage and AM printing stage, skipping the step1203of model transformations.

Aiming to balance the accuracy and efficiency, the proposed method integrated design stage and AM printing stage, skipping the step1203of model transformations. Figure 1 describes the framework of the proposed layer construction method, including four parts: a definition of layered model, information of the CAD model and AM process, a printing toolpath construction method, and an AM processing model. Here, we used this layered model instead of the traditional slicing model for toolpath generation. In the proposed method, a model information set that carries geometric constraints and design principles is reconstructed in the first step. Then, with the information of the original CAD model, the knowledge of AM (including AM parameters and AM constraints) and the predefined curved scanning patterns in the galvanometer scanning system, a printing toolpath construction method is used to generate the printing toolpaths.

Finally, an AM processing model, including printing toolpaths and AM parameters is constructed and sent to the SLM printer. The detail of the proposed method is described in the following.2.2. Definition of layered modelThe mathematical definition of a 2D shape is given by the following equation (Eq. (1)), and Figure 2 describes an example of a 2D shape. In the conventional AM data processing chain, a polygon is used to approximate the curved shape, which loses the accuracy of boundary and increases the filling time of the toolpath. Here, we use curves directly and predefine these curves in the galvanometer scanning system.Figure 2. An example of 2D shape.𝐺𝐺 = Ω𝑐𝑐 − Ω𝐻𝐻 = Ω𝑐𝑐 − �Ωℎ1 ∪ Ωℎ2 ∪ … ∪ Ωℎ𝑛𝑛�1204(1), where G is a 2D shape, surrounded by ith hole’s contour byis the domain surrounded by contour C, and (not includes the boundary hi).is the ith domain is a domain which is unitedΩ𝑐𝑐Ωℎ𝑖𝑖, andare given by Eqs. (2) and (3).

(8) defines a 3D layered model. Unlike the conventional continuous representation of the 3D model, the 3D layered model is a discretized model, which can be used in the generation of AM toolpaths directly.(8)𝑀𝑀 = {𝐿𝐿𝐺𝐺𝑖𝑖| 𝑖𝑖 = 1,2,3, … , 𝑛𝑛} , where M is the 3D layered model, layer, n is the number of total layers of the model.(as defined by Eq. (9)) is the layer graph model of the ith𝐿𝐿𝐺𝐺𝑖𝑖(9)𝐿𝐿𝐺𝐺𝑖𝑖 =< 𝐺𝐺𝑖𝑖, 𝑧𝑧𝑖𝑖 > , where can be calculated as Eq. (10):is the 2D shape in the ith layer,is the height of the ith layer. For SLM process,𝐺𝐺𝑖𝑖𝑧𝑧𝑖𝑖𝑧𝑧𝑖𝑖 (10)𝑧𝑧𝑖𝑖 = 𝑧𝑧1 + (𝑖𝑖 − 1) ∗ ℎ , whereis the height of the first layer, and h is the layer thickness.𝑧𝑧11205Figure 3. An example of a 3D layered model: (a) a STL format model. (b) a 3D layered model, 2D shapes of the (c) 10th layer, (d) 40th layer, (e) 70th layer, (f) 100th layer. (g) extruding curve for the first hole. (h) extruding curve for the second hole.2.3.

(b) a 3D layered model, 2D shapes of the (c) 10th layer, (d) 40th layer, (e) 70th layer, (f) 100th layer. (g) extruding curve for the first hole. (h) extruding curve for the second hole.2.3. Model information set constructionAs the STL format lacking the topological information of the model, it is unsuitable to use this format to represent the periodic cellular structure for slicing and toolpath planning. Referring to genes, they control the growth of animals and plants just by genetic information. Here, the model information set including the information of design space, constraints for layer graph and the design principles plays a role like a gene. However, there is no automatic method to construct the model information set because it needs human intelligence to extract the design principles according to the given model.

However, there is no automatic method to construct the model information set because it needs human intelligence to extract the design principles according to the given model. For a sample model, slicing is an efficient way to obtain a layer graph, however, for the periodic cellular structure with complex shape, slicing is time-consuming. Hence, the hull model is used in this paper to construct the geometric constraints of the layer graph instead of the original model. Figure 4 shows an example of the reconstruction of the geometric constraints and design principles.Figure 4. An example of the construction of the geometric constraints of layer graph and the design principles.Hence, the model information set can be given by Eq. (11).𝑀𝑀𝑆𝑆 =< 𝛺𝛺2𝐷𝐷, 𝐷𝐷𝐷𝐷 >1206(11), where the design principles.is the model information set,is a set of geometric constraints for layer, and DP is𝑀𝑀𝑆𝑆𝛺𝛺2𝐷𝐷2.4. AM information 2.4.1.

AM information 2.4.1. Predefined scanning pattern in galvanometer scanning systemExcept for the scanning patterns of line segment and polygon, the galvanometer offers a circular arc scanning pattern. Moreover, ellipse, B-spline and other curved scanning pattern are also available by predefining them in the controlling system of the galvanometer. Figure 5 presents the predefined circular arc scanning pattern in the galvanometer and the outputting format for the circular arc printing toolpath. Figure 6 presents the toolpaths and the corresponding scanning results using the predefined scanning patterns.Figure 5. (a) outputting format for circular arc printing toolpath; (b) schematic diagram of the predefinedcircular arc scanning pattern in the galvanometer.2.4.2. AM knowledgeIn the proposed method, AM knowledge includes AM parameters and AM constraints. It is mathematically defined as Eq. (12), more information of AM knowledge can be found in [14].

AM knowledgeIn the proposed method, AM knowledge includes AM parameters and AM constraints. It is mathematically defined as Eq. (12), more information of AM knowledge can be found in [14]. However, it is challenging to translate all the AM processing constraints to design principles or filters. This paper focuses on the layer construction method integrated the design and AM and does not discuss more the AM processing parameters and constraints.𝐴𝐴𝑀𝑀𝐴𝐴 = < 𝐴𝐴𝑀𝑀𝐴𝐴𝐴𝐴, 𝐴𝐴𝑀𝑀𝐴𝐴 > , where AMK is the AM knowledge, AMpa is a set of AM parameters; AMpa = {Mp: a set of material parameters, Pp: a set of processing parameters}, AMc is a set of AM constraints.1207(12)Figure 6. An example of printing toolpaths and the corresponding laser scanning results: (a) printing toolpaths; (b) the designed shapes; (c) the laser scanning results.2.5.

An example of printing toolpaths and the corresponding laser scanning results: (a) printing toolpaths; (b) the designed shapes; (c) the laser scanning results.2.5. Printing toolpath construction methodAs mentioned above, the conventional design & pre-processing method for AM process has limitations in balancing the accuracy and efficiency of the generation of printing toolpaths. Layer toolpath construction method [5] is a high-precision and high-efficiency toolpath generation method, connecting the design stage and printing stage for AM process. It divides the 3D design problem into many 2D design problems and generates toolpath directly. Figure 7 presents the workflow of the proposed printing toolpath construction method, and the main steps are described in the following:Step 1. Construct 2D design space. 2D design space is needed in the proposed method. It offers the geometric constraints for the 2D graphic design.

Construct 2D design space. 2D design space is needed in the proposed method. It offers the geometric constraints for the 2D graphic design. Designers obtain this 2D space by slicing the original 3D design space or even draw it directly for some simple shapes.Step 2. Design 2D layer graph. Design principles, predefined scanning patterns and AM processing constraints are considered to design the 2D shape. Unlike conventional interactive design mode, we use parameterized method to design the 2D shape by MATLAB or Rhino1208grasshopper. The type of the boundary is selected from the predefined scanning patterns in the galvanometer scanning system and a part of AM processing constraints are used in the size determinization, e.g., the cutting allowance. At this moment, only polygon, polyline, circle, circular arc, ellipse, and their combinations are allowed.

At this moment, only polygon, polyline, circle, circular arc, ellipse, and their combinations are allowed. Once the 2D layer graph is designed, another part of AM processing constraints, e.g., the overhang angle, are used as filters to check whether the 2D layer graph can meet the AM processing requirements.Step 3. Generate layer toolpath. The toolpath is a set of predefined scanning patterns in galvanometer scanning system, attaching AM processing information, i.e., laser power, scanning speed. A homemade clipping algorithm is used to clip the segment pattern toolpath of the parallel segments that are inside of the 2D shape.Step 4. Optimize layer toolpath. To reduce the printing time, the scanning directions of the segment patterns are set to ‘Zigzag’, and the scanning sequences of the different regions or contours are treated as a travel salesman problem (TSP), a genetic algorithm (GA) is used to solve this problem.Step 5. Add toolpaths and parameters in AM processing model.

Add toolpaths and parameters in AM processing model. The last step is to combine the toolpaths and other AM parameters, i.e., material parameters, AM processing parameters and add them in an AM processing model for outputting.Figure 7. Workflow of the proposed printing toolpath construction method.1209Case study3.1. Original modelTo demonstrate the method efficiency, a complex case, flame arrester, with periodic fine features from industry is selected. Figure 8 shows the original model of the flame arrester element. It has more than 10,000 holes and the STL model updating time is about 2 hours in a home computer (RAM: 4*8GB, Crucial, CPU: Xeon E5 2660 V4, Intel) with Rhino grasshopper.Figure 8. Original CAD model of the flame arrester element.3.2. Layer graphic model reconstruction 3.2.1. Model information set of the flame arresterAccording to Figure 3, many 2D shapes are designed to compose a 3D model instead of designing it directly.

Layer graphic model reconstruction 3.2.1. Model information set of the flame arresterAccording to Figure 3, many 2D shapes are designed to compose a 3D model instead of designing it directly. The 3D design space needs to be divided into 2D geometric constraints at a different height. The 2D geometric constraint of the flame arrester element at z = zi is given by Eq.

The 3D design space needs to be divided into 2D geometric constraints at a different height. The 2D geometric constraint of the flame arrester element at z = zi is given by Eq. (13).𝛺𝛺𝑅𝑅1 = Ω𝑐𝑐 − Ωℎ0 = {(𝑥𝑥, 𝑦𝑦)|𝑓𝑓𝐶𝐶(𝑥𝑥, 𝑦𝑦) ≤ 0} − {(𝑥𝑥, 𝑦𝑦)|𝑓𝑓ℎ0(𝑥𝑥, 𝑦𝑦) < 0} = {(𝑥𝑥, 𝑦𝑦)|𝑓𝑓𝐶𝐶(𝑥𝑥, 𝑦𝑦) ≤ 0, 𝑓𝑓ℎ0(𝑥𝑥, 𝑦𝑦) ≥ 0}𝐶𝐶: 𝑓𝑓𝐶𝐶(𝑥𝑥, 𝑦𝑦) = 𝑥𝑥2+ 𝑦𝑦22 − 75ℎ0: 𝑓𝑓ℎ0(𝑥𝑥, 𝑦𝑦) = 𝑥𝑥2+ 𝑦𝑦22 − 5.051210(13)(14)(15), where contour C, 𝛺𝛺𝑅𝑅1 Except for the geometric constraints, the original model also exhibits the designis the 2D design domain for the graph in the first layer,is a domain surrounded byis a domain surrounded by hole’s contour.Ω𝑐𝑐Ωℎ0ℎ0principles, the results are given as follows: (1) the circular holes are arranged at the vertices and the centers of hexagons in the hexagonal array; the hexagonal array at z = z1 is defined by Eq. (16), the circles can be represented by Eq. (22), and the hole’s region is given by Eq.

However, the printed sample will be cut from the substrate. It needs to design a cutting allowance for the flame arrester element. According to AM knowledge, the commonly used cutting allowance is 1mm and the layer thickness is 0.03mm. Hence, the 3D layered model of the flame arrester element can be achieved by Eq. (26):(26)𝑀𝑀 = {< 𝐺𝐺𝑖𝑖, 𝑧𝑧𝑖𝑖 >| 𝑖𝑖 = 1,2,3, … ,366}(27)𝑧𝑧𝑖𝑖 = 𝑧𝑧1 + (𝑖𝑖 − 1) ∗ 0.03(28)𝐺𝐺𝑖𝑖 = Ω𝑅𝑅𝑖𝑖 − Ω𝐻𝐻𝑖𝑖 ; Ω𝑅𝑅𝑖𝑖 ∈ Ω2𝐷𝐷; Ω𝐻𝐻𝑖𝑖 ∈ Ω𝐻𝐻2𝐷𝐷 , whereis the layered model of the flame arrester element,is the ith layer height, and, 𝑀𝑀 0.03𝑚𝑚𝑚𝑚 Figure 9 presents the first layer graph of the flame arrester element at the height of z = 0.03mm. As shown in Figure 9(a), the redesigned space for the flame arrester element is a cylinder with a hole. The geometric constraints for every layer keep the same, hence the geometric constraints of the flame arrester element can be given as follows:is the ith layer graph,is given by Eq. (16), and 𝑧𝑧𝑖𝑖is given by Eq.

ResultsFigure 10(a), (c) and (d) present the printing toolpaths of the first layer in the flame arrester elements. Figure 10(b) and (e) show the print samples. In the proposed method, all the filling toolpaths are selected from the predefined curved scanning patterns in the galvanometer. They are more accurate than the sliced contour of the STL model. Table 3 shows the comparison results of the time consumed by the layer construction method and traditional design & slicing method. The new method almost reduces 90% of the data processing time in AM processing chain.Table 3. The comparison results of the time consumed by the layer construction method and traditional design & slicing method.

Printing toolpaths and results: (a) toolpaths of the first layer in the Ø150mm *10mm flame arrester element, (b) printing result of the Ø150mm *10mm flame arrester element, (c) toolpaths of the first layer in the Ø50mm *10mm flame arrester element, (d) a description of the toolpaths , (e) printing result of the Ø50mm *10mm flame arrester element.ConclusionsThe paper proposed a novel layer construction method of periodic cellular structure for AM process, integrating the design stage and printing stage. It is much more efficient than the traditional design & slicing method. It has potential application in the high performance and multifunctional structure with complex geometries. However, there is still a lot of work that needs to be done in the future to improve this method, e.g., spline scanning pattern and efficient filling algorithm. Meanwhile, in the proposed method, the model information set is constructed according to the original model.

and Zhang, Y. (2020) "Design for additive manufacturing: Framework and methodology", CIRP Annals, Vol.69 No.2, pp: 578-599. [10] Liu, J. and To, A. C. (2017) "Deposition path planning-integrated structural topology optimization for 3D additive manufacturing subject to self-support constraint", Computer-Aided Design, Vol.91, pp: 27-45. [11] Wang, C., Qian, X., Gerstler, W. D. and Shubrooks, J. (2019) "Boundary Slope Control in Topology Optimization for Additive Manufacturing: For Self- Support and Surface Roughness", Journal of Manufacturing Science and Engineering-Transactions of the ASME, Vol.141: 091001-1. [12] Wang, Z., Zhang, Y., Bernard, A. (2021) A constructive solid geometry-based generative design method for additive manufacturing. Additive Manufacturing. Vol. 41:101952. [13] Yu-an, J., Yong, H., Jianzhong, F., Wenfeng, G. and Zhiwei, L.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEFFICIENT MULTISCALE PREDICTION OF CANTILEVER DISTORTION BY SELECTIVE LASER MELTINGC. Li, J. F. Liu, Y.B. Guo Dept. of Mechanical Engineering, The University of Alabama, Tuscaloosa, AL35487, USAAbstractLarge tensile residual stress is one major issue for metal components made by selective laser melting (SLM). Residual stress is induced by non-uniform heat input, which leads to part distortion and detrimentally affects product performance. The conventional single track simulation method is not feasible to predict the distortion of a macro part since it demands an exceedingly long computational time. The coupling multiphysics phenomenon during the SLM process further complicates this issue.

The coupling multiphysics phenomenon during the SLM process further complicates this issue. In this study, a temperature-thread multiscale modeling approach has been developed to predict part distortion of a twin cantilever. An equivalent body heat flux calculated from the micro scan model was imported as the “temperature-thread” to the subsequent layer hatch model. Then the hatched layer with temperature field can be used as a basic unit to build up the macro part. The temperature history and residual stress fields during the SLM process were predicted. And the distortion of twin cantilever was calculated with a reasonable accuracy compared to the experimental data.Keywords: Selective laser melting, distortion, multiscale simulation, additive manufacturing1. IntroductionSelective laser melting (SLM) is capable of producing near full density metallic functional component directly from CAD data [1,2].

IntroductionSelective laser melting (SLM) is capable of producing near full density metallic functional component directly from CAD data [1,2]. In a typical commercial SLM machine, the powder material is locally melted by a fast moving laser source (scan speed up to 15 m/s). The non- uniform heat input as well as the rapid heating and cooling in the material would generate large amount of tensile residual stresses in the component, which leads to part distortion and cracks [3,4]. Part distortion due to tensile residual stress is one of the major problems of SLM parts. It not only reduces the part geometrical accuracy but also and detrimentally affects the functional performance of the end-use parts.Several experimental works have been done to systematically investigate the formation of part distortion during SLM process. Kruth et al [5] studied the thermal deformation of a 1 mm thickness substrate with one layer of iron powder deposited.

Kruth et al [5] studied the thermal deformation of a 1 mm thickness substrate with one layer of iron powder deposited. Six different scanning strategies were taken into account, and island scanning strategy was proved to cause less distortion than that of the line scanning strategy. Kruth et al [6] identified the residual stress and part distortion of SLMed part using bridge curvature method. After the part was removed from the substrate, the part distortion in terms of angle α of the two bottom surface was measured to represent the Corresponding author Tel.: 1-205-348-2615; fax: +1-205-348-6419. E-mail address: yguo@eng.ua.edu236magnitude of thermal residual stress generated inside the component. Several parameters were studied including the length of scan vectors and rotation angle of scan vectors for two sequential layers. It has been found that a shorter scan vectors length and larger rotation angle tend to generate less deformation during the SLM process.

It has been found that a shorter scan vectors length and larger rotation angle tend to generate less deformation during the SLM process. Buchbinder et al [7] studied the method of reducing distortion of an aluminum twin cantilever by preheating the substrate (see Fig. 2). The cantilever which is an overhang structure can be distorted easily during the SLM process, thus a support structure was designed to minimize the deformation of the cantilever during the process. The cantilever distortion was measured after the support structure was removed from the substrate. Different cantilever bar thickness and preheating temperatures are investigated. A significant reduction of the cantilever distortion was observed when the substrate was preheated to 150 °C.To have a fundamental understanding of distortion mechanism for SLM processed part, numerical modeling as a powerful tool has been widely used to predict residual stress and part distortion in SLM.

Simulation works have been done on a small domain (usually single track or single wall). Dai et al. [8] simulated the effect of powder-to-solid transition to investigate the residual stress and distortion of metal and ceramic powders at the micro-scale. Hodge et al. [9] studied the thermal and mechanical history of a SLM process on the meso-scale (12 layers of powder). A volumetric moving flux was used to melt powder materials with material state change taken into consideration during the process. Aggarangsi et al. [10] explored the residual stress reduction method of a SLMed thin-wall structure by using a secondary moving heat flux to preheat the powder material in a finite element model. Heigel et al. [11] studied the thermal history and mechanical response of a single wall deposited on a single-side fixed substrate, the deflection history of the substrate was simulated and validated.Several studies predicted residual stress and part distortion in SLM on the macro part scale.

Zaeh et al. [12] applied a uniform thermal load adjusted from experimental data to heat up 20 real layers at the same time to predict the temperature and residual stress field of a cantilever with support on the substrate. Denlinger et al [13] applied a hybrid quiet inactive element activation strategy to perform 3D thermal-mechanical analysis of large parts on the order of meters. The distortion of the large part was validated with an electron beam deposited part. Prabhakar et al. [14] directly applied a uniform heat source simultaneously to the whole layer to simulate the residual stress formation of tensile test coupons produced by electron beam melting (EBM) as well as the distortion of the substrate.

Li et al [15-17] developed a stress-thread based multiscale simulation approach to fast predict part distortion, the temperature field of a melt pool was obtained through a single laser scan model and then mapped to a 5 mm by 5 mm small domain from which a local residual stress field was calculated and extended to a macro part model with different scanning strategies considered to predict distortion on a macro level, but only one layer was deposited on the substrate.Thermal stress prediction for several single tracks on a micro-scale using a fine mesh would take several hours to complete, and it depends on the model size and computer performance.

Moreover, a practical SLM part on the macro-scale is generally consist of millions of micro- scale laser tracks which dramatically increases the computational time for the coupled analysis.237Therefore, it is extremely difficult to predict part distortion of a practical SLM part if every single tracks is modeled even a powerful work station is applied.The objective of this study is to develop a temperature-thread multi-scale finite element approach for efficient predicting part distortion and residual stress for a practical SLM processed part by: (a) developing a novel equivalent heat source for a micro-scale single laser scan model; (b) modeling the power-liquid-solid material state transition phenomena during the SLM process; (c) predicting the residual stress field and distortion of a practical part on a macro-scale model; and (d) validating the predicted distortion with the experimental data.2.

Multiscale Simulation Methodology2.1 Simulation frameworkA multi-scale modeling approach (Fig. 1) is highly needed to predict practical part distortion and residual stress with low computational cost and acceptable accuracy. The temperature-thread multi-scale approach developed in this study can be described in three scales, i.e., micro-scale, meso-scale and macro-scale. In the micro-scale model, a single track is deposited using a laser scan and the temperature history of the melt pool is recorded. In the meso-scale model, the temperature history obtained from the micro-scale model is extended to a single layer. In the macro-scale model, a practical part is built layer upon layer by applying the thermal load developed from the meso-scale model.The three length scales are integrated through the temperature history “temperature-thread” of the melt material.

The temperature information is transferred from the micro-scale model to the macro-scale model using this multiscale method.Microscale laser scan model (Thermal)Moving 3D laser heat flux  Metal powder meltedOutputHeating-up time  Temperature historyMaterial statePowder-liquid-solidMesoscale layer hatch model (Thermal)Macroscale part model (Thermal-mechanical)Development of body heat flux  Heat up block layer simultaneouslyResidual stress field  Part distortionFig. 1 Multiscale modeling methodology.The temperature-thread method is described as follows: Firstly, the temperature field of a molten pool was calculated using a moving heat flux to melt powder material placed on the substrate in the micro-scale scan model. Thermal history on the center point of the scan vector is recorded. Secondly, an equivalent heat source is developed based on the thermal history in the238micro-scale scan model, and directly applied to the meso-scale hatch model.

Secondly, an equivalent heat source is developed based on the thermal history in the238micro-scale scan model, and directly applied to the meso-scale hatch model. Finally, the thermal history of one hatch layer is applied to the macro-scale part model, and each hatch layer is activated one by one until the whole part is built. The total computational time can be dramatically reduced using the multiscale approach compared to the conventional prediction methods.2.2 Material state transition mechanismsThe twin cantilever (see Fig. 2) was fabricated using an EOSINT M 270 SLM machine equipped with an Yb-fiber laser with a wavelength of 1064 nm. After the part was built, the support structure was removed from the substrate, and then the cantilever distortion was measured. A laser absorption ratio of 0.09 is used [18].

After the part was built, the support structure was removed from the substrate, and then the cantilever distortion was measured. A laser absorption ratio of 0.09 is used [18]. The process parameters are listed in Table 1.Table 1 SLM process parametersLaser power Scan speedW 195mm/s 800Scan spacing μm 150Layer thickness μm 30PartSupportSubstrate(a) Cantilever building strategyCutSubstrate(b) Support cut from the substrateSubstrate(c) Cantilever distortion after support removedFig. 2 Distortion of SLM parts: (a) cantilever building strategy, (b) support cut from the substrate, (c) cantilever distortion after support removed.The metal powder used in this study is a commercial available powder AlSi10Mg. Physical properties of AlSi10Mg are listed in Table 2 [19]. The temperature-dependent thermal material properties were used to ensure calculation accuracy, are listed in Table 3.

Physical properties of AlSi10Mg are listed in Table 2 [19]. The temperature-dependent thermal material properties were used to ensure calculation accuracy, are listed in Table 3. The plasticity of solid AlSi10Mg was obtained from literature [20], from which the AlSi10Mg tensile test samples were239fabricated by SLM process. The yield point of the material at 600 oC is assumed to 1/10 of that at room temperature. Plasticity properties are listed in Table 4.

The yield point of the material at 600 oC is assumed to 1/10 of that at room temperature. Plasticity properties are listed in Table 4. The substrate used in this study is an AlSi1MgMn plate.Table 2 Physical properties of AlSi10Mg [19]Elastic Modulus GPa 71Poisson's ratio0.33Density g/cm3 2.68Latent heatJ/kg 389000Melting point oC 600Table 3 Thermal properties of AlSi10MgTemperature (oC) Solid conductivity (W/mK) Specific heat (J/kg oC)20 110 739100 150 922200 159 797400 155 922Table 4 Plasticity of SLM processed solid AlSi10Mg [20]Temperature (oC) 20 20 600 600Stress (MPa) 200 300 20 30Plastic strain 0 0.01089 0 0.01089ABAQUS subroutine USDFLD was used to simulate the material state transition (powder- liquid-solid) during the SLM process. The subroutine is coded based on the SLM temperature history and powder’s melting temperature.

The subroutine is coded based on the SLM temperature history and powder’s melting temperature. Different properties for powder, liquid, and solid state materials were modeled based on the corresponding material state functions.2.3 Multiscale simulation modelsMicro scan model: In the micro-scale scan model, the temperature field of the melt pool was calculated using the commercial FEA package ABAQUS/Standard. A moving Gaussian distributed heat flux (see Fig. 3a) is developed to model the heat input of the scanning laser in the micro scan model. The power intensity of heat source is determined by laser absorption coefficient (cid:1827) of the powder material, laser power (cid:1842), laser spot size (cid:1856)(cid:3046) and the coordinates of the laser spot center (x, y). The thermal history of the center point of the scan track was recorded as shown in Fig. 3a.240Scan spacing (cid:1834)Body heatSpot dia.

The thermal history of the center point of the scan track was recorded as shown in Fig. 3a.240Scan spacing (cid:1834)Body heatSpot dia. (cid:1856)(cid:3046)vTemp.h t p e dlo o p(cid:3040) (cid:1856)t l e MSubstrateSubstrate(a) Micro scan model(b) Meso hatch modelTemp.Heat convection loss to atmosphereBlock layerSymmetric planePartHeat conducted to substrateSubstrateHeat loss to surrounding powderSupport structureResidual stress DistortionCutting planeBuild-up directionFixed bottom(c) Macro part modelFig. 3 Over-view of temperature-thread multiscale modeling approach.Meso Hatch Model: The shape of the melt pool was simplified to a cuboid with a length of laser spot diameter, a width of scan spacing, and a depth of melt pool. The equivalent body flux was defined as the power density (W/m3) which represents the input power for unit volume of melt material.

The equivalent body flux was defined as the power density (W/m3) which represents the input power for unit volume of melt material. Based on the thermal history of the melt pool calculated from the micro scan model, an equivalent heat source model was developed as shown in Fig. 3b. The total exposure time of the heat source in the meso hatch model is represented by the total heating time obtained in the micro scan model. The heat input was modeled as a body heat flux (cid:1869) which is associated with laser power (cid:1842), laser absorption coefficient (cid:1827), laser spot diameter (cid:1856)(cid:3046), melt pool depth (cid:1856)(cid:3040), and hatch spacing (cid:1834). The body heat flux q was given by Eq. (1):(cid:1869) (cid:3404)(cid:1827) ∙ (cid:1842) (cid:1856)(cid:3046) ∙ (cid:1856)(cid:3040) ∙ (cid:1834)Macro Part Model: One quarter of a twin cantilever with support structure was modeled as shown in Fig. 3c.

3c. The initial temperature of both the powder and substrate was set to 20 °C. The bottom surface of substrate was fixed during the SLM process. Three heat transfer mechanisms were modeled in the macro part model, i.e., heat conduction to the substrate, heat convection of the melt pool to the surrounding powder bed and atmosphere, and heat radiation to the atmosphere.In this study, the part along with the support was sliced into 12 layers in Y direction (part build-up direction as shown in Fig. 3c). At the very beginning of the simulation, all the elements in the part and support regions were deactivated. When the first powder layer was placed, the241(1)corresponding elements set was activated and heated up by the equivalent body flux for 0.4 milliseconds (exposure time calculated from the micro scan model). A cooling cycle for the whole model with 10 seconds was added right after the heating cycle. Then a new powder was added, heated up, and cooled down to room temperature.

A cooling cycle for the whole model with 10 seconds was added right after the heating cycle. Then a new powder was added, heated up, and cooled down to room temperature. This adding-heating-cooling cycle stopped until the whole part was built. At last, the elements located in the cutting plane were deactivated in the model to simulate the process of removing the support structure from the substrate. In order to predict the distortion of the cantilever after the SLM process, the cutting plane was removed by deactivating the associated elements in the plane to release the internal residual stress after the whole part including the support was build. A final distortion of the cantilever will be captured and compared to the experimental measurement.3. Model Prediction and Discussions3.1 Material state transitionThe material state contours and the corresponding temperature fields during the cantilever building process are shown in Fig. 4.

Model Prediction and Discussions3.1 Material state transitionThe material state contours and the corresponding temperature fields during the cantilever building process are shown in Fig. 4. The newly added powder material (in green) was placed on top surface of the cantilever and heated by the equivalent body heat flux. When the temperature went up and was higher than the melting point, the powder started to melt and the state of the material changed to liquid (in red). After the powder was heated for 0.4 milliseconds, the body heat flux was taken off and the material state is all in liquid. Then a cooling period which is ten seconds long was followed. During cooling, the liquid material began to solidify to solid material (in blue). Some of the solid material in the previous layer was remelted when heating up the next layer as shown in Fig. 4d. At last, the liquid was fully solidified and the whole model was cool down to room temperature.Fig.

4d. At last, the liquid was fully solidified and the whole model was cool down to room temperature.Fig. 4 Material state and temperature field evolution during layer build up.2423.2 Residual stressFig. 5 shows the predicted von Mises and three normal residual stress contours of the cantilever before and after the support structure was released from the substrate. With the support on the substrate, the maximum von Mises stress (approximately 300 MPa) was found on the top layer of the cantilever arm and the connection region between the cantilever and the substrate. And it is higher than the yield strength (200 MPa) which means plastic deformation occurs after the SLM process.

And it is higher than the yield strength (200 MPa) which means plastic deformation occurs after the SLM process. After the support was cut from the substrate, the maximum von Mises stress decreased to approximately 150 MPa, stress concentration was observed at the connect corner of substrate and cantilever.von Mises Stressvon Mises StressS11S11S22S22S33S33(a) With support on substrate(b) Support released from substrateFig. 5 Residual stress contour of part model: (a) with support on substrate, (b) support released from substrate.243With support on the substrate, a large tensile residual stress S11 was observed on the upper layer of the cantilever as shown in Fig. 5a. Also, the normal stress S11 which is perpendicular to the part building direction, dominates the three normal stress components. This could be explained by the temperature gradient mechanism [5].

5a. Also, the normal stress S11 which is perpendicular to the part building direction, dominates the three normal stress components. This could be explained by the temperature gradient mechanism [5]. During the heating cycle, the expansion of the top layer of the cantilever was restricted by its surrounding material and compressive stress was induced. When the top layer cools down, the contraction of the top layer was then constrained, which leads to tensile residual stress on the top layer of the cantilever. After the support was cut from the substrate, normal stress (S11) in X direction on the top layer of the substrate becomes compressive. It was caused by the plastic deformation of the cantilever arm after equilibrium residual stress state was broken inside the as-build cantilever.3.3 Part distortionIn the experiment, the measurement points for distortion were located on the top surface of the twin cantilever along longitudinal direction.

To compare with the measured distortion, a nodal path located on the top surface of the cantilever along the distortion direction (Z) was created in the simulation shown as the black dots in Fig. 6.2.41.9)m m( n oi t r ot s Di1.40.9ExperimentPrediction0.40.10510152025 Distance (mm)3035404550Fig. 6 Part distortion validation: prediction vs measurement.After the support was removed from the substrate, the cantilever bends towards the build direction, a concave up shape curve was observed. The predicted distortion and the experimental data show same bending direction and similar magnitude of distortion. The formation of this curve is due to the thermal history of the twin cantilever. The material located in the upper layer of the cantilever was expanded due to the laser heating.

The formation of this curve is due to the thermal history of the twin cantilever. The material located in the upper layer of the cantilever was expanded due to the laser heating. When the material cools down, the plastic strain in the upper layers became smaller than the lower layers, and a large tensile residual stress was induced in the upper layer areas. After the support was removed, most of the tensile residual stress was released and the cantilever reached a stable state. Then, a concave shaped distortion was formed. The maximum distortion in the experimental was 2.1mm while24455the predicted maximum distortion was 1.5 mm, which means an error of about 28 % was obtained. One possible reason could be the input heat source underestimated the actual laser heat input during the SLM experiment. Furthermore, the number of sliced layers has been reduced from approximately 400 in the experiment to 12 in the macro part model.

Furthermore, the number of sliced layers has been reduced from approximately 400 in the experiment to 12 in the macro part model. If more sliced layers were used in the simulation, one could expect more accurate predictions.4. ConclusionsA temperature-thread efficient modeling approach has been developed for fast prediction of a cantilever distortion produced by SLM. Thermal information has been transferred through micro- scale laser scanning, meso-scale layer hatching, and macro-scale part build-up. The predicted distortion of the twin cantilever was compared with the experimental data.

The predicted distortion of the twin cantilever was compared with the experimental data. The key findings of this study are summarized as follows:An equivalent heat source has been developed from the thermal history of melt pool in micro-scale laser scan model and applied to the meso-scale hatch layer, which is incorporated in the macro-scale part to predict part distortion.The material state transition phenomenon (powder-liquid-solid) during the SLM process was modeled using different field functions based on the temperature history, maximum temperature and melting temperature of the material.Large tensile residual stress was found on the top layer of the part with support on the substrate. After the support was removed, tensile residual stress in the upper layer of the cantilever decreased by 70%.Same distortion trend with a maximum distortion error of 28% was predicted when compared to the experimental data.5. References[1]J. Kruth, M. Leu, T.

26(1):1-10.[8] K. Dai, L. Shaw, 2004, Thermal and mechanical finite element modeling of laser formingfrom metal and ceramic powders, Acta Materialia. 52(1):69-80.[9] N. Hodge, R. Ferencz, J. Solberg, 2013, Implementation of a thermomechanical model indiablo for the simulation of selective laser melting, .[10] Aggarangsi P, Beuth JL, 2006, Localized preheating approaches for reducing residualstress in additive manufacturing. In Proc. SFF Symp 709-720.[11] J. C. Heigel, P. Michaleris, E. W. Reutzel, 2015, Thermo-mechanical model development and validation of directed energy deposition additive manufacturing of Ti–6Al–4V, Additive Manufacturing. 5(0):9-19.[12] M. F. Zaeh, G. Branner, 2010, Investigations on residual stresses and deformations inselective laser melting, Production Engineering. 4(1):35-45.[13] E. R. Denlinger, J. Irwin, P. Michaleris, 2014, Thermomechanical modeling of additive and Journalmanufacturing Engineering.

Two variations of the algorithm have been implemented and tested on a Farsoon HT251P machine using (polyamid) PA3300 polymer powder. The algorithm is based on building a statistical predictive model of the objective response (i.e., maximization of tensile strength), aggregating the constraint function (i.e., limited amount of warping), in an iterative manner by simultaneously improving the accuracy of the predictive model as well as searching for the optimum set of process parameters. The difference in two algorithmic variations is the number of samples to update at each iteration. While the first method is based on a single sample update, the latter searches for multiple simultaneous updates to let the manufacturer try several potentially good sets of parameters in the same machine to eventually speed up the experimental evaluation procedure.IntroductionAdditive manufacturing refers to a set of processes that deposit layers of material in orderto form an object.

One such popular technique is Selective Laser Sintering (SLS) that primarily works by depositing and melting layers of powders. Objects of great complexity can be produced from a computer aided design (CAD) file, making this very popular for their ability to quickly fabricate prototypes and parts that would be time intensive to construct using conventional methods.The mechanical properties of printed parts are an important consideration for bothprototyping and manufacturing. The part must meet certain mechanical and physical requirements for use or testing. Material properties such as particle size distribution and reflectiveness are an important consideration [1], but they are not the only factors affecting the mechanical properties of the printed part. Those properties are also dependant on process parameters such as laser power, scan spacing, and scan speed [2].

Those properties are also dependant on process parameters such as laser power, scan spacing, and scan speed [2]. Finding the optimum process parameters that produce the highest quality parts is critical.951Mathematical modeling of the process may seem like a possible way to obtain theoptimum parameters. Pharhami and McMeeking create a system of equations whose solutions predict the movement of particles during the sintering process [3]. Zhou et al. present a model that simulates the part bed upon which heat transfer analysis needs to be done [4]. While the results of these models appear to be accurate, they require significant computational time and power due to the large number of elements. Additionally, deriving the formulas is a complicated process which may not generalize well to different materials.Singh and Prakash also create a mathematical model of the SLS process to predict density[5].

However, Singh and Prakash used a factorial design of experiments approach to collect data used to predict their equation for density. Factorial design of experiments (DOE) is an experimental method for obtaining information by combinatorially varying controllable factors. Ghanekar and Crawford. use this method to characterise the SLS process for finding the parameters that produce optimal parts [6]. However, the cost of exploring the parameter space using factorial DOE increases exponentially with the number of parameters. To address this, Singh and Parkash add an initial screening stage to reduce the number of parameters, making it a two factorial design [5]. Liu et al. explore the parameter space using Taguchi’s method [7]. The number of experiments is reduced by selecting a small set from all the possible combinations of factors based on an orthogonal array. Negi et al.

The number of experiments is reduced by selecting a small set from all the possible combinations of factors based on an orthogonal array. Negi et al. use a face-centered central composite design to reduce trials [8].Because SLS parameters are actually continuous, results from a DOE process can then be fitted to a mathematical formula for better analysis and optimization. Negi et al. and Sachdeva et al. (2013) fit their results using a quadratic model [8, 9]. This allows them to discover optimums that exist between the values chosen for a DOE design. Ning et al.

Negi et al. and Sachdeva et al. (2013) fit their results using a quadratic model [8, 9]. This allows them to discover optimums that exist between the values chosen for a DOE design. Ning et al. use the collected data to train a neural network to predict the optimum process parameters [10].In this paper we propose a constrained efficient (single/parallel infill) sampling-based optimization algorithm to find the optimum SLS process parameters for improving the tensile strength of a tensile bar while reducing the number of physical experiments.MethodologyExperiments were carried out on a Farsoon HT251P SLS-machine using PA3300polymer powder ( (optimization) parameters. Tensile strength of the printed parts is aimed to be maximized subject to the limit based on product quality (i.e., warped tensile specimens are considered to be infeasible solutions).Fig. 1). The laser power and scan spacing are considered to be the control see ​​952Figure 1.

1). The laser power and scan spacing are considered to be the control see ​​952Figure 1. (Left) Farsoon HT251P SLS-machine used for the experiments. (Right) An image captured during the scanning process.Fig. 2 shows how the width and height measurements are considered in order to evaluatethe performance of each process parameter setting. Each experiment consists of six tensile test parts printed at the specified parameters. The dimensions of the part are measured before the tensile test is performed. Afterwards, the tests with the highest and lowest tensile strength are discarded. The remaining four results are then averaged to determine the final tensile strength.Figure 2. A tensile bar and how the width and height measurements are taken shown schematically.953The initial experiments are chosen using the optimal latin hypercube sampling (LHS),which is a uniform space filling sampling method.

The same set of initial experiments was used for both methods.Laser Power (W)Scan Spacing (mm)Tensile Strength (MPa)250.1744.9530.345.3320.2627.7460.2148.5390.08NaN600.12NaNTable 1. Initial experiments and their observed tensile strength. Results for 39W/0.08mm and 60W/0.12mm were unable to be obtained (NaN).Experiments at 39W/0.08mm and 60W/0.12mm showed significant warping. The resulting tensile strengths are inaccurate and most of the printed parts could not be tested at all. The objective value of the tensile stress measurement is instead considered to be,f = Current worst tensile strength - total deformation * 1.5where total deformation is the difference between the average dimensions of viable test parts from the dimensions of the warped one.

We are aware of the fact that tensile strength and deformation have different units (i.e., MPa and mm, respectively), but our intention is to assign a fitness value, to an infeasible solution, lower than the current worst acceptable (not warped) solution. This approach, turning a constrained problem into an unconstrained one, will eventually guide the search towards better regions of the design space.Overall efficient sampling framework is shown in Fig. 3. First, the optimal LHS producesthe initial sample set of six tensile bars for building the first surrogate (y) of the objective function (i.e., tensile stress of the printed specimen). Each physical experiment considered to be an expensive or time consuming design evaluation, Y.954(1)Figure 3. Surrogate based optimization frameworkDue to the nonlinear correlation between various part properties and machine parameters,Gaussian Process (GP) or Kriging is chosen to construct the surrogate.

In addition to being superior to a polynomial response surface model or artificial neural networks (in case of limited number of samples) for approximating complex nonlinear functions, the process of constructing a GP surrogate also provides an estimate of the mean square error (MSE). These output of the surrogate, i.e., y and MSE, are then used to calculate the expected improvement (EI) function, which is a statistical acquisition function to be maximized to find new update points to improve the fitness value upon the current best process setting found so far. The EI is given below in Eq. (2),where Φ and φ are the cumulative distribution function and probability density function of a normal distribution, respectively, s(x) is the MSE estimate at new candidate sample vector x, and y is the best observed value so far. The maximum of the EI surface is sought using a genetic ​best algorithm. In this paper we investigated using a singular infill method and a parallel infill variation.

The maximum of the EI surface is sought using a genetic ​best algorithm. In this paper we investigated using a singular infill method and a parallel infill variation. More details about the computation of EI can be found in [11]. This methodology was previously applied in pultrusion and friction stir welding processes [13, 14], respectively. Next two subsection give more details about the update strategy, i.e. the number of sample points (optimal candidates).​955(2)1. Singular InfillThe first method will return a single infill point per iteration of the algorithm. The fitnessof a point is defined as its EI value.2. Parallel InfillThe second method returnsis defined by the user. To k infill points per iteration, where k ​ ​​​prevent the infill points from being clustered tightly around one optimum, a points fitness is modified depending on the closeness of its neighbors. We use a shared fitness formulation, Eq.

We use a shared fitness formulation, Eq. (3), proposed by Goldberg [12],(3)and σ j and i where d( ) is the distance between points j ​ ​ ​ the diversity or the number of niches. If the distance between two points is greater than σ then they will not affect each other's fitness value. The fitness value, , of the samples locating f’ ​ ). This will f in more crowded regions will be lower than their original fitness value ( ​ automatically let the algorithm locate a number of diverse and optimal solutions in a single optimization run., i ​is a predetermined value controlling​​​​​share​, ​share​​​However, clustering is possible even after the transformation. To address this, the geneticalgorithm is modified to account for distance between optima.

To address this, the geneticalgorithm is modified to account for distance between optima. In order to be considered to be part of the top the set and also be at least a certain distance apart from every member of the set.set of infill points, a point must have a better fitness than at least one member of k ​​ResultsThe first set of tests used the singular infill method. Five infills were tested with one newsample gained per day.956Infill #Laser Power (W)Scan Spacing (mm)Tensile Strength (MPa)160.00.2547.9252.40.2146.3343.80.1844.3448.00.2246.3543.00.242.1Table 2. Singular Infill Results⚫Feasible Initial Experiment ⚫Infeasible Initial Experiment ◼​​​Figure 4. Surrogate constructed after five infill points.Infill Point ​Infill Point ​The singular infill algorithm searches the space around the initial experiment with thehighest tensile value.

Surrogate constructed after five infill points.Infill Point ​Infill Point ​The singular infill algorithm searches the space around the initial experiment with thehighest tensile value. Low experimental results around that point has led to a high fitness estimate in the top right corner (i.e, high power and scan spacing).957Infill #Laser Power (W)Scan Spacing (mm)Tensile Strength (MPa)160.00.2546.0252.40.2146.4343.80.1845.4448.00.2246.0543.00.247.6Table 3. Singular Infill Confirmation ResultsThe confirmation results show less variation than the experimental results. This leads usto believe that there may be outside factors affecting the experiment.Experiment #Laser Power (W)Scan Spacing (mm)Tensile Strength (MPa)156.80.2240.4159.40.339.5149.00.2345.3242.10.1743.2245.20.1844.5247.90.1744.2Table 4.

Parallel Infill Results958⚫Feasible Initial Experiment ⚫Infeasible Initial Experiment ◼​​​Figure 5. Surrogate constructed after six infill points.Infill Point ​Infill Point ​The parallel infill method exhibits similar search behavior. The first experiment explored the top right corner (i.e, high power and scan spacing) and it no longer exhibits the high extrapolation visible in Fig.4.Experiment #Laser Power (W)Scan Spacing (mm)Tensile Strength (MPa)156.80.2243.0159.40.350.3149.00.2348.5242.10.1748.9245.20.1846.3247.90.1748.7Table 5. Parallel Infill Confirmation Results959Discussion and Further WorkIn this section, the results of the optimization algorithm as well as improvements thatcould be made in the future are discussed. After each method was completed, a confirmation test was run on the infill points to examine variation in the previous iterative results.

After each method was completed, a confirmation test was run on the infill points to examine variation in the previous iterative results. Differences in tensile strength can be observed between both graphs. These differences might be due to the use of recycled powder. We believe more consistent results can be obtained if experiments are done on a larger scale with greater control over other factors.Both singular and parallel infill methods appear to quickly converge to an area near thecenter of the graphs (i.e., medium power and scan spacing), which matches operator expectations for parameters that lead to high tensile strength. Both infill methods also produced similar surrogates, but parallel infill methodology was able to achieve this in two days while singular needed five. It must be noted that areas in the surrogate with few or no observations are extrapolations.While the initial stage has yielded promising results, future improvements can bepursued.

It must be noted that areas in the surrogate with few or no observations are extrapolations.While the initial stage has yielded promising results, future improvements can bepursued. Variation in observed results must be accounted for, and existing literature in mathematical optimization propose a variety of methods that may be used. Since mechanical properties are determined by more than two parameters, future experiments should be done to confirm the validity of these methods with more parameters. Additionally, future work can build upon this framework to implement optimization for multiple mechanical properties.Conclusion This paper presents an efficient sampling framework for optimizing Selective LaserSintering process parameters to improve the tensile strength of the printed product. The mechanical properties of the printed parts are important in both manufacturing and prototyping.

The mechanical properties of the printed parts are important in both manufacturing and prototyping. It is known that a variety of process parameters affect the final product, however the exact relationship is hard to quantify. Physical experiments to determine the optimal process parameters are typically done using factorial design of experiments methods which are both expensive and time consuming. As the number of process parameters tested grows, the number of experiments needed increases exponentially. In this paper two variations of an efficient sampling method to algorithmically determine optimal parameters has been proposed. Both employ a surrogate function to estimate the true relationship between parameters and the target property. The difference between the two variations is the number of experiments per iteration of the algorithm.

The difference between the two variations is the number of experiments per iteration of the algorithm. The first method is based on a single sample update while the second searches for multiple possible samples.960The methods were tested with a Farsoon HT251P machine using (polyamid) PA3300 polymer powder. Laser power and scan spacing were optimised for tensile strength. Not only singular and parallel variations showed similar performance, but also the parallel variation required less time to find the same optimal solution. While the algorithms showed similar optima that matched operator experience, further improvements can be made to account for variation in test results and optimising for multiple mechanical properties.AcknowledgementsThis research was conducted with the generous support of Farsoon Technologies.References[1] Y.A Song, W.

ficient Three Dimensional Modelling of Additive Manufactured Textile StructuresG. A. Bingham*Loughborough Design School, Loughborough University, Loughborough, UK, LE11 3TUAbstractTextile structures realised by Additive Manufacturing (AM) techniques have received increasing attention during the previous decade. Due to their potential to significantly improve upon both the geometric complexity and functionality available from conventional fibre-based textiles, AM textiles present a serious opportunity to design and develop novel solutions for conventional and high-performance textile applications.

AM textiles also provide the capability to produce net-shape textile artefacts and allow the development of personalised, high-performance textiles from a variety of materials currently being processed by AM technologies.While the motivation exists for the wider-scale adoption of these novel structures, practical access to an efficient three dimensional (3D) modelling strategy limits their applications. The research presented here discusses the issues surrounding the 3D modelling of complex AM textiles and discusses dedicated methodologies developed for the generation of their conformal data.

The research presented here discusses the issues surrounding the 3D modelling of complex AM textiles and discusses dedicated methodologies developed for the generation of their conformal data. The research culminates with a robust methodology for the generation of AM textile apparel data suitable for manufacture by AM techniques.IntroductionResearch in the area of Additive Manufactured (AM) textiles has mainly concentrated on the development of efficient modelling strategies for the creation of the three dimensional (3D) Computer Aided Design (CAD) data required for their manufacture by AM techniques [Bingham et al 2007, Bingham 2007]. However, additional research has also been undertaken to address their mechanical properties, design and possible applications of these novel, complex structures [Crookston et al 2008, Johnson et al 2011].

Most recently, research has been undertaken to successfully design and manufacture AM textiles capable of attaining the British level one standard for stab-resistance [Johnson et al 2012]. While a genuine motivation exists for the further development of AM textile applications, efficient modelling of conformal and net shape AM textile artefacts remains a significant issue.Modelling AM textiles with CADThe modelling of individual AM textile link structures using currently available 3D CAD systems is relatively simple for expert users and complex geometric designs can be generated with relative ease. The use of CAD systems can be expanded further for the generation of planar samples of AM textile using a simple planar ‘Array’ or ‘Patterning’ function, common to most current CAD systems.

The use of CAD systems can be expanded further for the generation of planar samples of AM textile using a simple planar ‘Array’ or ‘Patterning’ function, common to most current CAD systems. Planar samples can be generated as either quadrilateral-based or triangular-based patterned arrays as demonstrated in Figure 1.260Figure 1: AM textile samples, Left - quadrilateral-based, right - triangular-basedThe limitations associated with current CAD systems relate to the capability of generating conformal or net-shape AM textile artefacts. While it is possible to generate simple net-shape AM textiles (e.g.

While it is possible to generate simple net-shape AM textiles (e.g. cylindrical), the generation of more complex net-shapes, such as hemispherical samples, is problematic and can result in incorrect tessellation of the individual link structures, also demonstrated in Figure 2.Figure 2: Cylindrical and hemispherical AM textile samplesFurther investigation revealed the generation of a hemispherical AM textile sample could only be achieved through the use of a triangular-based AM textile link configuration (instead of a quadrilateral-based) and the acceptance of dimensional variance within the individual link structures. The modelling of the hemispherical AM textile sample demonstrated in Figure 2 also required a significant amount of manual manipulation by the CAD user and was therefore laborious, error prone and inefficient.

While this can be tolerated for smaller net-shaped AM textile artefacts, the modelling of larger AM textile structures, for example apparel, would be impractical.In reality, an individual textile link structure can be designed with an extremely high level of geometric complexity. However, this initial experimental modelling demonstrates the difficulty encountered when attempting to combine these links into conformal or net-shape AM textile structures. In this case, the requirement is not the creation of a singular or linear and circular array of complex individual geometries, but the automated creation of a uniformly distributed collection of complex geometries, potentially numbering in the thousands, that ultimately create the hierarchy of a conformal AM textile structure.

As complex (double curvature) geometries cannot be accurately described by any simple array or patterning function within current CAD systems and no automated mapping function currently exists to replace the required manual manipulation of individual link geometries, the generation of conformal AM textile data is currently severely restricted.

The work presented here has developed a novel approach to reduce these restrictions, allowing complex AM textile structures to be modelled quickly and efficiently.261AM textile modelling strategyIn order to unlock the potential of conformal AM textiles and provide practical access to the efficient generation of AM textile data, a dedicated modelling strategy is required, with specific constraints:Provide an efficient means of generating complex conformal AM textile data suitable formanufacture by AM techniquesAttempt to maintain uniformity within the dimensions of the individual AM textile linkstructuresRemove the need for manual intervention by the user for the positioning of the individualAM textile link structuresUniformity within the final AM textile structure is desirable due to manufacturing constraints and the potential relationship between geometry (size, shape) and functionality textile structures [Bingham 2007].

Major when considering high-performance AM dimensional variation within the individual AM textile links could result in either a failure to build (the links being too small, highly deformed or touching another link) or failure to perform as designed (scale, shape and functionality). Observations from experimental modelling using current CAD systems has highlighted the efficiency of using an array or patterning function to describe the intended locations of individual link geometries and repeating the original link geometry to those intended locations, as demonstrated in Figure 3, creating the final AM textile structure.Figure 3: Array functioning (quadrilateral-based)An efficient modelling strategy can therefore be defined as an approach that removes the requirement of any manual intervention of the AM textile link structures by the CAD user and accurately controls their location and spacing by automated mapping of the individual geometries.

Repeating the original link structure also removes any potential for dimensional variance within the resultant AM textile. The intended locations of the individual links (shown as dashed lines) in Figure 3 can be considered as a surface mesh, representing the desired final configuration of the conformal or net-shape AM textile. This insight developed the idea of generating and utilising a ‘mapping’ surface mesh for the generation of conformal and net-shaped AM textiles.

This insight developed the idea of generating and utilising a ‘mapping’ surface mesh for the generation of conformal and net-shaped AM textiles. However, while the strategy of generating a surface mapping mesh is relatively simplistic for planar and cylindrical samples as demonstrated in Figure 3, the generation of a mapping mesh for complex topology is much more problematic.Surface mapping meshThe main requirement of the surface mapping mesh, when used as a means of providing locations over complex topology for the accurate positioning of AM textile link geometries, was a uniform structure and equidistant node spacing, demonstrated in Figure 4.262Figure 4: Idealised mapping meshFor a surface mesh to exhibit such qualities, the only possible structure the mesh can be generated from, is a set of equilateral triangular elements.

To clarify, if a section of a uniform and equidistant surface mesh is considered, as demonstrated in Figure 4, the structure of the mesh can be described through a system of nodes and elements. The nodes of the mesh, shown as red spheres in Figure 4, are locations from the intended topology that are described by the mesh. The elements of the mesh, shown as grey cylindrical geometries in Figure 4, are described as the geometry connecting the nodes of the mesh, completing the surface mesh structure. For the surface mesh to be uniform, each node contained within the mesh must be connected to a maximum of six other nodes in a triangular, hexagonal configuration.

For the surface mesh to be uniform, each node contained within the mesh must be connected to a maximum of six other nodes in a triangular, hexagonal configuration. Also, for the mesh to be equidistant, the distance observed between each node, shown as (x) in Figure 4 must all be constant, and therefore, the elements of the mesh must all be equal length.Various existing surface mesh generation algorithms were investigated for their application as a mapping mesh generator for AM textile data generation, including: Indirect or parametric space types [Marur 2005], Advancing front types [Borouchaki 2000], Direct and indirect sphere packing types [Shimada 1997]. Additionally, several commercially available Finite Element (FE) pre-processors were also explored, including: Fluent Gambit [Fluent 2011], MSC Patran [MSC 2011], ANSY [ANSY 2011] and Hypermesh [Hypermesh 2011].

However, it was quickly established that the generation of a completely uniform and equidistant mesh structure as demonstrated in Figure 5 was not possible. Modification of an existing surface mesh generation algorithm was also considered, however, given the exacting requirements of the uniform and equidistant mesh structure, it was decided that a dedicated surface meshing algorithm was required.Further research was undertaken to achieve this aim and a new experimental surface meshing algorithm based on Sphere Packing [Shimada 1997] was developed by the author specifically for AM textile data generation [Bingham 2007]. The results of the research are demonstrated in Figure 5 for a range of topology, showing the desired uniform and tolerance- based (typically +/- 0.2units), equidistant mesh structure required when utilised as a mapping mesh for AM textile data generation.

Having established a methodology for the generation of a suitable mapping surface mesh investigation into a robust geometry mapping methodology was required.263Figure 5: Example uniform and equidistant surface meshesDevelopment of a mapping tool for AM textile data generationThe initial 3D modelling of AM textile structures highlighted many limitations when generating conformal or net-shaped AM textiles and the requirement for a dedicated 3D modelling strategy or tool. However, it also reinforced the geometric complexity achievable when considering the design and 3D modelling of AM textile link structures and the efficiency of mapping to a predefined mesh. A requirement of any specifically developed modelling strategy for the generation of AM textile data must also provide this level of geometric complexity for the design and development of present and future AM textiles.

It was therefore decided to incorporate the capabilities of 3D CAD systems into the modelling strategy by simply importing the 3D data they generate. Similarly, it was decided not to directly incorporate the surface meshing algorithm discussed previously into the methodology but supply the necessary data as a further input. These decisions simplify the requirements of a dedicated tool for the generation of conformal AM textile data into a system capable of mapping complex 3D data to an imported surface mapping mesh.

These decisions simplify the requirements of a dedicated tool for the generation of conformal AM textile data into a system capable of mapping complex 3D data to an imported surface mapping mesh. The mapping methodology can be summarised in the following four stages:Stage one: Import CAD data of link geometry (input one) and surface mesh data oftopology (input two)Stage two: Calculate surface normals at mapping locations from surface mesh Stage three: Orientate link geometry to match mapping locations of the surface mesh Stage four: Translate orientated link to mapping locationThe four individual stages of the mapping methodology provide the core functionality of the developed AM textile modelling strategy. The entire AM textile modelling strategy can be summarised as follows:1. Design individual AM textile link structure using conventional CAD software and export as STL data2. Identify target surface topology for the final AM textile structure 3.

Design individual AM textile link structure using conventional CAD software and export as STL data2. Identify target surface topology for the final AM textile structure 3. Generate uniform and equidistant mapping mesh of target surface topology 4. Identify mapping locations (node and or element midpoints) and calculate surface normal from mapping mesh5. Copy and orientate original AM textile link structure to match required surface normal at the mapping location while remaining at the global origin6. Translate orientated AM textile link structure to mapping location 7. Repeat stages five and six until every mapping location has been utilised 8.

Translate orientated AM textile link structure to mapping location 7. Repeat stages five and six until every mapping location has been utilised 8. Export final AM textile as STL data while omitting original AM textile link structure264The core mathematics underpinning this methodology is not discussed within this paper but is fully discussed in [Bingham 2007, Bingham 2012].Experimental testing of the mapping methodologyTo validate the capabilities of the documented methodology it was coded into a discrete software mapping tool [Nottingham 2006].

The software mapping tool has three key functions that required testing and validating:The ability to calculate the surface normal at the mesh nodes and element midpoints The ability to map complex geometry data described in STL format to the specifiedmapping locations without distortion or user interventionThe ability to maintain the original rotational orientation of the mapped geometrySuccessfully attaining these capabilities would therefore provide all the functionality required to efficiently generate conformal AM textile data.Experiment one: Surface normal validationThe surface normal at any mesh node or element midpoint was calculated directly from the surface mesh and not the original geometry from which the mesh was created. To understand the effectiveness of the mapping tool in delivering this key capability, marker geometry, as demonstrated in Figure 6, was mapped to example surface meshes demonstrated in Figure 5.

To understand the effectiveness of the mapping tool in delivering this key capability, marker geometry, as demonstrated in Figure 6, was mapped to example surface meshes demonstrated in Figure 5. The marker geometry was designed to allow a visual inspection of the final generated data to be undertaken as the direction of the mapped arrow indicated the surface normal at each mapping location. Any anomaly with the surface normal calculation at the mapping locations was then easily detected through the visual inspection of the arrow directions of the resultant STL data.

Any anomaly with the surface normal calculation at the mapping locations was then easily detected through the visual inspection of the arrow directions of the resultant STL data. The mapping meshes, demonstrated in Figure 5, used for this experiment were:Planar mapping mesh: 60 units square with a 6 unit nodal spacing Domed (hemisphere approximation) mapping mesh: 60 units square with a 6 unit nodalspacingSaddle mapping mesh: 60 units square with a 6 unit nodal spacingFigure 6: Marker geometry265Experiment one resultThe marker geometry demonstrated in Figure 6 was mapped to both the nodes and element midpoints of the three example surface meshes. The resultant STL data generated was then imported into 3DS Max [Autodesk 2012] to visually inspect the positions of the marker geometry relative to the mapping locations and the surface normal calculated at each location.

The results of nodal mapping for each of the three surface meshes are demonstrated in Figure 7.Figure 7: Experiment one, marker geometry nodal mappingFigure 7 demonstrates that the orientation of the marker geometries reflect the expected surface normal at the nodal mapping locations of the three meshes utilised. For each mesh, the mapping tool positioned the marker geometry to the intended mapping locations (nodes and element midpoints) while also matching the expected surface normal, resulting in a smooth transition of arrow direction and reflecting the curvature of the intended surface geometry.Experiment two: Complex geometry mapping validationThe second experiment was conducted to determine whether the mapping methodology was capable of mapping complex geometry without any distortion. Complex test geometry (Mobius knot) was utilised for this experiment as demonstrated in Figure 8.

Complex test geometry (Mobius knot) was utilised for this experiment as demonstrated in Figure 8. The test geometry, described as STL data, was mapped to each of the surface meshes detailed in experiment one (planar, dome and saddle), firstly to the nodes of the mapping mesh and then to the element midpoints.Figure 8: Experiment two test geometry, mobius knotExperiment two resultThe results of the mapping experiment to the nodes of the example domed mapping mesh are demonstrated in Figure 9. For all three surface meshes utilised, the mapping methodology was successful in mapping the geometry to the nodes and element midpoints of the mesh by modifying the orientation to the required surface normal without distorting the266original geometry.

The individual shell entities within the generated STL data were extracted and compared with the original input STL data, where no discernible differences were recorded.Figure 9: Experiment two nodal mapping to domed meshExperiment three: Rotational orientation validationThe third experiment was conducted to determine whether the original rotational orientation of the mapping geometry could be maintained within the final mapped geometries. Three asymmetrical test geometries were utilised for this experiment as demonstrated in Figure 10.

Although identical, the rotational orientation (Cartesian axes) for each one was modified to produce a different final outcome:Teapot one: Spout orientated to the X axis Teapot two: Spout orientated to the Y axis Teapot three: Spout orientated to a combination of the X and Y axisFigure 10: Experiment three test geometry orientations, teapotThe test geometries, described as STL data, were mapped to each of the surface meshes detailed in experiment one and two (planar, dome and saddle meshes), firstly to the nodes of the mapping mesh and then to the element midpoints.Experiment three resultThe results of the mapping experiment to the nodes of the example domed mappingmesh for all three test geometries are demonstrated in Figure 11, 12 and 13.267Figure 11: Experiment three, test geometry one, nodal mapping to domed meshFigure 12: Experiment three, test geometry two, nodal mapping to domed meshFigure 13: Experiment three, test geometry three, nodal mapping to domed meshFor all three surface meshes utilised, the mapping methodology was successfully in mapping the geometry to the nodes and element midpoints of the mesh by modifying the orientation to the required surface normal without distorting the original geometry.

A visual inspection of the generated data for each test geometry indicated that the mapping methodology was also successful in maintaining the original rotational orientation of the test geometries throughout the mapping process.Discussion of experimental resultsThe results of the three validation experiments indicate that the mapping methodology developed has the capability to efficiently map complex geometry described as STL data to a surface mesh without any distortion, whilst also matching orientation to the surface normal at each location. These capabilities therefore provide all the functionality required for the efficient generation of conformal AM textile data. Through the application of a uniform and equidistant mapping mesh it is possible to efficiently generate complex and conformal AM268textile data suitable for manufacture by AM techniques.

Through the application of a uniform and equidistant mapping mesh it is possible to efficiently generate complex and conformal AM268textile data suitable for manufacture by AM techniques. The results of this process are demonstrated in Figure 14, where a 1unit spacing mapping mesh based on a 60 unit square saddle surface has been utilised for the generation of a conformal AM textile sample. The modelling process for this sample required no user intervention other than the input of the original textile link structure (simple torus) as STL data and the surface mesh data file. The final STL data as demonstrated in Figure 14 was generated in under 30 seconds.Figure 14: Conformal AM textile data generation using the developed mapping methodologyTo validate the suitability of the data generated for manufacture by AM, the STL data was inspected within Materialise’s Magics software [Materialise 2012].

The complete AM textile STL data structure was first modified so each individual link of the AM textile became a discrete part using the ‘shells to parts’ functionality of the software. Once complete, an interference check was performed between all parts to identify any interfering triangles of the STL data. The inspection revealed no interference within the entire structure and therefore suitable for manufacture by AM.Limitations of the methodologyThe demonstrated methodology, for the generation of conformal AM textile data, is robust, accurate and efficient. However, the quality and range of the data generated is dependent on the input mapping mesh, where any distortion or irregularity within the mapping mesh is directly replicated in the resultant data.

However, the quality and range of the data generated is dependent on the input mapping mesh, where any distortion or irregularity within the mapping mesh is directly replicated in the resultant data. The capabilities of the experimental surface meshing algorithm developed for this research are currently limited and can only generate uniform and equidistant (tolerance-based) mesh structures from mathematically described quadratic surface patches. The surface patches are described using five-point polynomial interpolation curves (splines) that form the four boundaries of each quadratic patch. The system requires the user to input five point (height) values for each side of the quadratic patch in order to generate a target topology (surface) that allows the creation of a uniform and equidistant mapping mesh structure.

Additionally, the experimental meshing algorithm generates the uniform and equidistant mesh structure by ‘mesh seeding’ from the centre of the target surface and prevents full mesh coverage of the entire quadratic patch, as demonstrated in Figure 5 and 14. While the experimental meshing algorithm allowed the discussed mapping methodology to be developed and validated, the limitations of the meshing algorithm prevents practical access to a range of more complex topology and manifold objects that are required for present and future applications of AM textiles. Addressing this limitation required a second investigation of high-quality surface mesh generation to be undertaken.269Mesh generation: Re-meshingThe initial investigation of surface meshing revealed no satisfactory solution for the generation of a uniform and equidistant mapping mesh structure.

A secondary line of investigation examined aspects of re-meshing that are normally utilised for the quality improvement of existing mesh structures for a range of applications, including; STL quality, FEA and rendering applications in polygon modelling systems. Re-meshing involves the manipulation of nodes and elements within an existing mesh structure based on a set of quality variables, typically element lengths and angle or aspect ratio. Normally these variables are defined within a maximum and minimum range and cannot be guaranteed globally throughout the resultant mesh structure, dependant on the underlying topology. Two excellent examples of Re-meshing software include Geomagics [Geomagics 2012] and Meshlab [Meshlab 2012], the latter being an open source shareware.

Two excellent examples of Re-meshing software include Geomagics [Geomagics 2012] and Meshlab [Meshlab 2012], the latter being an open source shareware. Examples of the mesh quality delivered by both systems are demonstrated in Figure 15 based on the original STL mesh structure of a hand.Figure 15: Re-meshing of hand STL data, middle right, Geomagics re-mesh and far right, Meshlab re-meshThe documented re-meshed examples are high-quality in comparison to conventional surface meshing algorithms but do not attain the quality expectations required for AM textile generation, where uniformity (essential) with limited dimensional variability is required. To utilise such meshes, the mapping methodology required a scaling function to be included that allows the link geometry to be modified (scaled) based on the element lengths at each node location within the mesh structure.

This was eventually included within the mapping software and the results of using the Meshlab re-mesh are demonstrated in Figure 16.Figure 16: AM textile glove using Meshlab re-meshing process270The use of a scaling function within the mapping methodology and re-meshed surface meshes does allow the efficient generation of net-shaped and manifold AM textile artefacts. However, the irregularity within the mapping mesh structure (node to element ratio) reduces the range of possible link designs that can be utilised due to mesh structure not being completely uniform. The use of re-meshing also removes the ability to control the dimensions of the resultant links or their separation. These limitations can result in the data not being suitable for manufacture by AM techniques due to the specific separation tolerances requirements of built structures, typically 0.3mm for Laser Sintering.

For some AM textile applications, the issues associated with using re-meshing can be tolerated, however, the aim of this research was the development of a method for the generation of uniform and dimensionally regular AM textile data.Mesh generation: Mesh conformingA further area of investigation was the notion of conforming a predefined mesh structure to a targeted surface topology. Mesh conforming has several qualities that make it an attractive alternative for the generation of AM textile data over re-meshing-based alternatives. The process of mesh conforming is demonstrated in Figure 15 and requires an initial planar mesh structure to be created with defined dimensions that is fitted (conformed) to a target topology based on specified variables, including mesh shear, element stretch and element compression [Autodesk 2012].

Control over these variables affects the defined planar mesh’s ability to conform to the targeted topology but also affects the final distortion observed in the resultant mesh structure (element lengths). Through careful management of these variables a compromise can be generated, resulting in a surface mesh that approximates the targeted topology with a completely uniform structure and contains minimal dimensional variation.Figure 17: Mesh conforming processThe application of mesh conforming in combination with the new scaling function within the mapping software allows conformal AM textiles data to be efficiently generated, as demonstrated in Figure 18.

This technique allows practical access to a range of complex topology far beyond the capabilities of the experimental meshing algorithm and produces higher-quality mesh structures (complete uniformity, typical element length +/- 0.5 units), and therefore higher quality AM textile data, then those generated by re-meshing techniques.271Figure 18: AM textile mask using mesh conforming processDue to the uniformity of the generated mapping mesh, this technique also allows a greater range of AM textile link designs to be utilised (not shown).

However, the use of mesh fitting for the generation of a mapping mesh does preclude the ability to generate net-shaped AM textile artefacts – this can only be achieved through re-meshing techniques as it is impossible to accurately conform a planar mesh to a manifold object without significant distortion at the boundaries of the fitted (conformed) mesh and the introduction of seams.AM textile apparel modellingA sub-aim of this research was the development of a methodology that would allow personalised (conformal) AM textile apparel data to be efficiently generated with suitable accuracy for manufacture by AM techniques. Using anthropometric data, or 3D body scanning data, the aim was the generation of a suitable mapping mesh that would allow accurate mapping of AM textile link structures for the production of a complete AM textile garment.

Initially it was hoped that re-meshing techniques would provide the required mapping mesh but the discussed limitations prevent its application. The alternative to re- meshing was the application of mesh conforming but this introduces the requirement to have discrete seams between individual panels of conformal AM textiles. While not fully desirable, seams are a recognised and necessary component of conventional garment design/construction and therefore it is appropriate that they also become a necessary component of conformal AM textile garments design/construction. The methodology required by the application of mesh conforming for the generation of personalised AM textile apparel data can be summarised as follows:1. Design individual AM textile link structure using conventional CAD software and export as STL data Identify target surface topology for the final AM textile apparel (3D body scan data or anthropometric data)1.

Design individual AM textile link structure using conventional CAD software and export as STL data Identify target surface topology for the final AM textile apparel (3D body scan data or anthropometric data)3. Design panel layout of AM textile apparel based on target surface topology from stage two4. Generate planar mapping meshes based on dimensional requirements of AM textile link structure from stage one and identified panels from stage three5. Mesh conform the planar mapping meshes to all indentified panels from stage three 6. Extract conformed meshes from stage five as a single mapping mesh and trim any overlapping mesh elements Import AM textile link structure STL data and final mapping mesh into mapping software5. Mesh conform the planar mapping meshes to all indentified panels from stage three 6.

Mesh conform the planar mapping meshes to all indentified panels from stage three 6. Extract conformed meshes from stage five as a single mapping mesh and trim any overlapping mesh elements Import AM textile link structure STL data and final mapping mesh into mapping software2728.Identify all mapping locations (node and/or element midpoints) and calculate surface normal from mapping mesh9. Copy original AM textile link structure and orientate to match required surface normal at the mapping location while remaining at the global origin10. Scale AM textile link structure based on the elements lengths at mapping location 11. Translate orientated AM textile link structure to mapping location 12. Repeat stages nine, ten and eleven until every mapping location has been utilised 13.

Translate orientated AM textile link structure to mapping location 12. Repeat stages nine, ten and eleven until every mapping location has been utilised 13. Export final AM textile apparel as STL data while omitting original AM textile link structureStages two, three and six of this methodology are illustrated in Figure 19.Figure 19: Target topology (left), panel layout (middle) and complete mapping mesh (right)The discussed methodology allows the generation of personalised AM textile garment data to be accurately generated, which is suitable for manufacture by AM techniques. However, the seams generated are a result of the interference between the individual panels of AM textiles and not a fully designed feature of the final AM textile garment as demonstrated in Figure 20, right.

Further work is required to provide a dedicated seam generation technique.Figure 20: Seams generated between individual panels of mapped AM textile link structuresThe complete AM textile garment data generated using this methodology is demonstrated in Figure 21.273Figure 21: Complete AM textile garment using mesh conforming processTo validate the suitability of the complete AM textile garment data for manufacture by AM techniques (Figure 21), the STL data was inspected within Materialise’s Magics software [Materialise 2012]. The complete AM garment was first modified so each individual link of the AM textile became a discrete part using the ‘shells to parts’ functionality of the software. Once complete, an interference check was performed between all parts to identify any interfering triangles of the STL data. The inspection revealed no interference within the structure except, as anticipated, for the seams.

The inspection revealed no interference within the structure except, as anticipated, for the seams. The interfering links at the seams where combined using the ‘Boolean’ functionality. The combined links at the seams introduce an element of solidity into the AM textile garment but still retain a level of movement.ConclusionThe aim of this research was to investigate and develop an efficient means of generating conformal AM textile STL data suitable for manufacture. The research presented here demonstrates the issues surrounding the 3D modelling of conformal AM textiles using conventional 3D CAD systems and presents a methodology for the efficient generation of such complex 3D data based on the application of mapping meshes. The conclusions of the research can be summarised as follows:AM textiles provide a real opportunity for the design and manufacture of geometrically complex and potential high-performance textiles structures.

However, practical access to an efficient means of generating conformal and net-shape 3D data restricts wider scale adoption and further investigation.274Current CAD software solutions do not have the required capabilities to generate conformal AM textile data but do provide the modelling capabilities required for the design of complex individual link structures.Mapping complex geometries to a predefined surface mesh is a highly efficient means of positioning large numbers of individual elements without any user (manual) intervention.

The presented methodologies provide the capability to map complex geometry described as STL data to an input surface mesh without any distortion of the mapped geometries or any user intervention.The presented methodology allows the original rotational orientation of the input geometryto be maintained throughout the mapping process.The limitations of the geometry mapping methodology for the generation of uniform and equidistant AM textile structures are directly related to the input mapping mesh. However, where uniformity and equidistance is not a requirement, the mapping methodology can be applied to generic surface mesh data from Re-meshing systems, FE pre-processors and polygon-based modellers for the generation of complex hierarchal structures of mapped geometries.Re-meshing techniques offer a viable solution for the generation of manifold mapping meshes but the quality of the data generated is not yet suitable for fully uniform and equidistant AM textiles.

The inclusion of a scaling function within the mapping software does address some of the issues discussed but the lack of uniformity restricts the range of AM textile links designs that can be utilised.Mesh conforming provides completely uniform but dimensionally variable mapping meshes for an extensive range of complex topology. This technique provides high-quality mapping meshes and extends the capabilities of conformal AM textile generation. However, issues surrounding boundaries and seams require further investigation.Further limitations relate specifically to the datasets generated using the mapping methodology. STL format was chosen to simplify the 3D data import and export process of the mapped geometries, and while this succeeded in simplifying the process, STL data is considerably larger than alternative CAD-based formats, for example STEP or Parasolid.

Therefore, the final dataset generated is a function of the imported data multiplied by the number of nodes and element midpoints utilised. For simple cases the data is entirely manageable, for larger surface meshes with thousands of nodes and element midpoints, combined with a complex mapping geometry, datasets can easily reach into the Gigabytes and therefore restricts the possibilities.Further workThe presented methodology allows the efficient generation of conformal AM textile data. However, the suitability of the data generated is directly dependent on the quality of the input surface mesh. An experimental surface meshing algorithm was specifically developed for the task of AM textile data generation but further work is required to extend its capabilities to more complex surface topology.

Alternatives for mapping mesh generation include re-meshing techniques and mesh conforming - further work addressing both of these techniques is required to extend the capabilities.The methodology for generating AM textile apparel data is effective but further work is required to simplify the process and provide an alternative for seam generation. Further work is also required to explore the design and application of AM textile possibilities that this new methodology enables.275ReferencesANSYS software website, http://ansys.com/, accessed September 2006Autodesk 3D Studio Max software website, http://www.autodesk.co.uk/adsk/servlet/pc/index?siteID=452932&id=14596087, accessed June 2012Bingham, G.A., Hague, R.J.M., Tuck, C.J., Long, A.C., Crookston, J.J.

and Hague, R.J.M.,“Finite Element Modelling of Mechanical Behaviour of Rapid Manufactured Textiles”. Proceedings of the Institution of Mechanical Engineers, Part L: Journal of Materials: Design and Applications, v 222, n 1, 2008, p 29-36Fluent software website, http://www.ansys.com/Products/Simulation+Technology/Fluid+Dynamics/ANSYS+FLUENT, accessed July 2011Geomagics, software website, http://www.geomagic.com/en/, accessed June 2012Hypermesh software website, http://www.altair.com/software/hw_hm.htm , accessed August 2011Johnson, A.J., Bingham, G.A. and Wimpenny, D., “Additive Manufactured Textiles for High-Performance Stab resistant Applications” – 12th RDPM conference, Lancaster, England, June 2011Johnson, A.J. and Bingham, G.A., “Additive Manufactured Textiles for High-Performance Stab Resistant Applications".

and Bingham, G.A., “Additive Manufactured Textiles for High-Performance Stab Resistant Applications". Submitted to the International journal of Rapid prototyping, September 2011 – Accepted, awaiting publicationMaterialise Magics software website, http://www.materialise.com/Magics, accessed March 2012Marur, S., “On the meshing of trimmed 3D surfaces”, Advances in Engineering Software, v36, pp 338–345, 2005)Mesh lab, software website, http://www.meshlab.org/, accessed June 2012MSC software website, http://www.mscsoftware.com/Products/CAE-Tools/Patran.aspx, accessed July 2011Nottingham, Textile Composites Research Group at the University of Nottingham website, http://www.textiles.nottingham.ac.uk/, accessed August 2011Shimada, K., Yamada, A. and Itoh, T., “Anisotropic triangular meshing of parametric surfaces via close packing of ellipsoidal bubbles”, Proceedings of the 6th International Meshing Roundtable, Park City, pp 375-390, 1997276.

forts to Reduce Part Bed Thermal Gradients during Laser Sintering ProcessingMengqi Yuan* and David Bourell*Laboratory for Freeform Fabrication, Mechanical Engineering Department The University of Texas at Austin, TX 78712AbstractPart bed surface thermal gradients (x-y plane) are usually present in laser sintering (LS) fabricators. The purpose of this study was to investigate various means to reduce these thermal gradients. Several experiments were conducted using a FLIR™ infrared camera to examine the thermal profile of the part bed during the LS operation. Experiments included thermal profile characterization of the part bed with different nitrogen shielding gas flow rates, assessment of the proper experimental settings, and a temperature profile record of the part bed from the warm-up to the cool-down stage.

A series of experiments were conducted using the laser as a heat source to preheat part bed surface cold spots to decrease the thermal gradients, which effect was limited by the natural low thermal conductivity of nylon 12 powder and large heat convection. Moreover, manifolds were mounted below the piston to provide warm nitrogen down draft flow during the LS operation.IntroductionLaser sintering is an additive manufacturing technology that uses a high power laser to fuse polymer powder into a mass that has a desired three-dimensional shape. The laser selectively scans and fuses powder material on the surface of the powder bed based on a previously generated CAD file. After one layer is scanned, the powder bed is lowered by one layer thickness, creating a new layer that is scanned. The process is repeated until the part is completed [1]. This process has the potential to be a valuable industrial tool [2].

The process is repeated until the part is completed [1]. This process has the potential to be a valuable industrial tool [2]. Thermal processing during the LS operation is essential in improving the efficiency of the process and the accuracy of product quality and dimension.Polyamide 12, also known as Nylon 12, is a thermoplastic material that is widely used in laser sintering. It has high elongation, good abrasion resistance, good specific strength, and melts at a temperature around 180°C [1]. Thermal management and heat transfer analysis of polyamide 12 in the LS part bed are essential to better understand the effect of processing on the service part properties and performance.Temperature distribution of powder bed during LS has been analyzed in different materials including metal, polymer, etc. Some were focused on the surface temperature distribution, while some others were about the LS operation chamber. S.

Some were focused on the surface temperature distribution, while some others were about the LS operation chamber. S. Kolossov et.al [3] developed a 3D thermal model of LS using finite element simulation, which incorporated the nonlinear behavior of the titanium powder. They found the temperature profiles during LS operation along both X- and Y- axes reached a maximum around 2500°C in the middle of the axes, and decreased to 500- 750°C on both sides.

They found the temperature profiles during LS operation along both X- and Y- axes reached a maximum around 2500°C in the middle of the axes, and decreased to 500- 750°C on both sides. These results were in agreement with the infrared red camera experiments.962Moreover, the only temperature profile examined was during the stable heat-up period; it is unknown how the part bed temperature profile behaved during LS processing which is an important part of the overall thermal exposure.There is a relation that illustrates the laser energy density with the laser power and beamproperties:ρ = P DV(1)where ρ is laser energy density (LaserED), P is laser power, D is laser beam diameter, and V is the scan speed [3].S. Kumpaty, D. Cottrill et al. [4] analyzed the heat transfer process in the Sinterstation® 2500Plus during the part build, recording the changes in Duraform® polyamide powder temperature using an infrared thermocouple.

[4] analyzed the heat transfer process in the Sinterstation® 2500Plus during the part build, recording the changes in Duraform® polyamide powder temperature using an infrared thermocouple. First, with the part bed at 149°C, tests were conducted on three square discs with 7,9,11 Watts of laser power. It was found the part bed temperature increased with increasing laser power. Second, the temperature profile of a single layer was explored. A valley was created on the temperature profile when cooler powder was spread over the part bed (140-150 °C). Temperature increased in the sintering process (170- 175 °C), and then the part bed powder cooled and solidified and the temperature stabilized around 161°C until the roller deposited a fresh layer of powder over the top surface. Last, a relation between the laser power and the temperature rise was found:T = 1.6111+ 0.6761P (2)where T is the temperature rise from 6 to 9 °C, P is the laser power from 6 to 12 Watts. Eq.

Last, a relation between the laser power and the temperature rise was found:T = 1.6111+ 0.6761P (2)where T is the temperature rise from 6 to 9 °C, P is the laser power from 6 to 12 Watts. Eq. (2) shows the temperature rise is proportional to the power increment, which is constant with Eq. (1), which shows the laser energy density is proportional to the power. However, the value of this prediction is limited due to the limited power range, and the thermocouple was found accumulated with powder at the end of the experiment. In addition, the sensor measurement circle is relatively small comparing to the part bed, which led to the experiment inaccuracy.Franco, M. Lanzetta and L. Romoli [5] conducted experiments to assess the energetic aspects of LS using polyamide 12 powder, and they analyzed the energetic productivity and laser parameters. A SYNARD CO2 laser was used in 3D LS machine.

A SYNARD CO2 laser was used in 3D LS machine. The energy density was varied modulating the scan speed and the laser power in the range 500-2500 mm/s and 12.5-50 W, respectively. The laser spot diameter was kept constant at 0.7 mm. A combination of 16 different groups of parameters was tested, resulting an energy density between 0.0071- 0.2857 J/mm2. Moreover, the sintered parts with different energy densities were attached at the surface of a container, which had a vibrating device beneath to measure the energetic intensity (EI). EI was found to be low ranging from 0.2- 0.6 J/mm3 with respect to an ED ranging from 0.015- 0.3 J/mm2. In addition, it was found that the best results could be obtained operating with energy density between 0.02-0.08 J/mm2.

In addition, it was found that the best results could be obtained operating with energy density between 0.02-0.08 J/mm2. Specially, the process with 0.06-0.08 J/mm2 laser ED could lead to elimination of preheating phase during the LS process, since most energy (35%) was consumed by the chamber heaters.963In this research effort, a 2D thermal profile investigation on the polyamide 12 powder bed during LS operation was conducted. The effect of shielding gas flow rate and laser pre-scan treatment were investigated as means to reduce thermal gradients on the surface of the part bed during LS processing.Experimental Procedure1. SS 2500 parametersIt is important to acquire proper operating temperature on left/right feed bins and the part bed.

SS 2500 parametersIt is important to acquire proper operating temperature on left/right feed bins and the part bed. There are basically two steps before the LS: heating to initial warm-up set-points (no powder is spread during this time); heating to operation set-points (0.1mm powder layer is spread over the surface, 25.4mm in total).To evaluate the most proper operating temperature for the feed bins, after reaching the initial warm-up set points on feed bins (around 70 °C), feed bins set-points were manually increased by 1- 2°C after every 2-3 layers of powder were added. The maximum feed temperatures were reached when the feed bed started to show subtle cracking, or ripples were observed on the part bed surface. The proper feed heater set-points were 1 °C below the maximum feed temperature.

The proper feed heater set-points were 1 °C below the maximum feed temperature. Moreover, the set-points might be different on left and right feed bins, due to the individual heater performance and heat leakage on each side.To evaluate the proper operating temperature for the part bed surface, careful observation of the heat fence part building is essential. First, the part bed temperature was set to a relatively low operating temperature, ~171°C). If slight curling was seen during a part build, the part bed set- point was increased by 2°C. The process was repeated until non-curling flat layers were obtained.2. Thermal profile investigationApparatus Measurements of the temperature at the surface of the part bed were made by means of an infrared (IR) camera, which was a FLIR model A325, serial 48000675 thermal imaging camera, with a 320×240 pixel resolution， 16-bit resolution, and focal length 18 mm. . The sensor in the camera measured long-wave (8-12 μm) IR radiation.

. The sensor in the camera measured long-wave (8-12 μm) IR radiation. For the measurement, the camera temperature range was 0-500 °C, yielding a thermal resolution of 0.03 °C. Absolute accuracy of the camera was 2 °C for known emissivity. Object emissivity was 0.825 by comparing the camera-reported temperature of a heated black body source with and without the SS 2500 front window. . Image acquisition rate was adjustable from 0-60 Hz.The insulated viewing window in front of the SS2500 was replaced with an insulated aluminum-fiberglass panel with visible and IR transparent zinc-selenide (ZnSe) viewports. The IR camera was mounted at the front of the machine and looked through the viewports. The ZnSe window had an anti-reflective coating for a transmissivity greater than 0.98 [6]. The camera was enclosed in a fan-cooled box to protect from dust and powder. The camera was controlled by FLIR ExaminIR software.

The camera was enclosed in a fan-cooled box to protect from dust and powder. The camera was controlled by FLIR ExaminIR software. Figure 1 shows a representative thermograph from the IR camera of a heated chamber, with the nonlinear temperature scale in °C.964Figure 1: Build chamber in the SinterStation 2500 surface that including heater, IR sensor,part bed and two feed bins, and IR thermograph during warm-up stage.Procedure To obtain a temperature profile of the part bed from heat-up to cool-down, which would last for several hours depending on the product size, the IR camera mounted on the ZnSe window was set with a data acquisition frequency of 0.5Hz. Data at every stage were recorded and exported to an Excel file and processed by MATLAB. For the warm-up stage, the part bed temperature increased after a new layer of powder was spread, and the maximum value was selected in each layer to perform a smooth plot of the part bed mean temperature.

For the LS stage, the transient equilibrium temperature of the sintered part after sintering was selected and plotted to show the smooth mean temperature profile on sintered parts at different positions of the part bed. For the cool-down stage, the transient equilibrium temperature after each layer was selected and plotted.To analyze the circulating nitrogen thermal effect on the part bed surface temperature, the part bed was heated to 175 °C. First the nitrogen flow rate was changed from 0.5 m3/h to 2.5 m3/h, increased by 0.5 m3/h incrementally and held for 5 minutes. A line in the middle of the part bed along the X- axis was observed by the IR camera to obtain a thermal profile. After that, the circulating nitrogen valve was closed and the line was observed by the IR camera for 1 hour. Thermal profiles were plotted along the X- axis at different time periods (0 second after nitrogen valve closed, 800 seconds later, 2400 seconds later, and 3600 seconds later).3.

Thermal profiles were plotted along the X- axis at different time periods (0 second after nitrogen valve closed, 800 seconds later, 2400 seconds later, and 3600 seconds later).3. Laser pre-scan treatmentOne approach to achieving a more uniform part bed surface temperature is to pre-scan cooler areas with the laser spot. Referring to Eq. (1), the parameters affecting the laser energy density are the laser beam diameter, laser scan speed and laser power. The laser beam diameter is difficult to change “on the fly”, and the change of scan speed is limited with SinterStation functions. The laser scan power is easier to change and was selected as the assessment parameter.965Tests were conducted on 15 cubes (0.0254m× 0.0254m × 0.0254m) with laser power varying from 0.5-7 watts. The part bed temperature was held constant at 173 °C. The space between each square along the x- and y-axes was 0.00635m.

The part bed temperature was held constant at 173 °C. The space between each square along the x- and y-axes was 0.00635m. The IR camera was mounted during the LS stage to acquire the temperature rise after the sintering scan with data acquisition frequency 2 Hz and default emissivity 0.825. The temporal duration of the temperature rise was also recorded. Graphs presenting the temperature change versus power were plotted and analyzed.Moreover, to evaluate different powers and position effects on the products, mechanical properties of sintered tensile bars were tested using an INSTRON® Series 3340 Model 3345 Single Column Testing System, with a maximum capacity of 5kN and maximum vertical test space of 1123 mm [7]. Experimental data were analyzed using INSTRON Bluehill Lite software. The extension rate was set at 2 mm/min with a sensitivity of 40%.Results and Discussion1.

Experimental data were analyzed using INSTRON Bluehill Lite software. The extension rate was set at 2 mm/min with a sensitivity of 40%.Results and Discussion1. SS 2500 parametersThe SinterStation left and right feed bin and part bed operation temperatures varies by machine and heater. For the machine and heater used in this case, the left feed bin operation temperature was 100 °C, while the temperature for the right feed bin was 93 °C.Part bed operation temperature was set at 173-175°C after several observations and experiments. Previously using a lower set point led to curling on the sintered part. Too high a setpoint melted or glazed the powder surface.2. Thermal profile investigationTemperature changes at every stage (warm-up, LS and cool-down) were recorded by the IR camera and processed by MATLAB.

Thermal profile investigationTemperature changes at every stage (warm-up, LS and cool-down) were recorded by the IR camera and processed by MATLAB. Figures 2a and 2b show the actual and maximum temperature profile versus time during the warm-up stage.Figure 2a shows that the temperature dropped 10-20 °C each time the roller spread over the polymer surface and then increased by 1-2 °C compared to last cycle maximum temperature. The maximum temperature plateaued at 175-178 °C. Figure 2b simplified the warm-up process by disregarding the inverse powder-spreading spike. The temperature increased steadily from 148 °C to 172 °C in about 1000 seconds then slowly reached an equilibrium around 178 °C which held for 2700 seconds. The equilibrium temperature starting at 2000 seconds was quite close to the polyamide 12 melting point of about 180 °C [1].An IR image of LS on the powder bed surface is shown in Fig. 3.

The equilibrium temperature starting at 2000 seconds was quite close to the polyamide 12 melting point of about 180 °C [1].An IR image of LS on the powder bed surface is shown in Fig. 3. There was a 10-15 °C temperature difference on the powder bed surface between the hottest spot and coldest spot. The heater is shown all white because its temperature falls outside the measurement range. The part bed surface is hot in the middle and right corner, but cold around four sides and the upper-left corner. Convection from the environment may contribute to the thermal gradient.966Figure 2: Polyamide 12 part bed surface temperature profile versus time in warm-up stagebefore LS. (a) Part bed surface mean temperature plotted by IR camera and (b) Maximum mean temperature as processed by MATLAB.Figure 3: Thermograph from the IR camera during laser sintering. Default emissivity is set as 0.825. The temperature range was manually set to be 120- 200 °C.

Default emissivity is set as 0.825. The temperature range was manually set to be 120- 200 °C. Two squares drawn on the surface (lower right and upper left) show a hot spot and cold spot of powder bed. The bright yellow rectangle in the middle of the part bed surface is a square that just finished being laser sintered. The brighter white rectangular spot on right was being laser sintered when the image was captured.967Figure 4a-c show powder bed surface temperature profiles during laser sintering that wererecorded, exported by FLIR camera and processed by MATLAB.Figure 4: Polyamide 12 part bed surface temperature profile versus time during lasersintering.

(a) The mean temperature of part bed surface during LS; (b) The mean temperature profile of hot spot surface during LS; (c) The mean temperature profile processed by MATLAB that includes the mean surface temperature profiles versus time for the part bed, hot spot and cold spot.Figure 4a shows a thermal recording from the IR camera versus time for the part bed surface mean temperature during LS. A large temperature gap occurred after each 70-80 seconds indicating the moment the roller spread the powder over the surface, while the mean temperature dropped to 152-153 °C.Figure 4b is the real-time IR record showing the hot spot mean temperature on the part bed surface versus time. The high temperatures shown (300-400 °C) were not reasonable since the part was in a melt or half-melt status, and the emissivity was changed, rendering the measurement to be inaccurate.

The high temperatures shown (300-400 °C) were not reasonable since the part was in a melt or half-melt status, and the emissivity was changed, rendering the measurement to be inaccurate. However, it still seen that the LS process temperature was decreased by the roller spreading powder over the surface and was increased by LS.968Figure 4c shows the mean temperature profile versus time processed by MATLAB of the whole part bed, hot spot, and cold spot during the whole LS process. Each data point was selected at the moment after LS when the temperature reached equilibrium. However, the temperature data for the hot spot and cold spot were higher than 180 °C because the part was still in a “half melt” status with the emissivity changed from 0.825. In addition, the system error was included (2 °C) [6]. It is seen that the hot spot temperature was 3-5 °C higher than that of the cold spot.

In addition, the system error was included (2 °C) [6]. It is seen that the hot spot temperature was 3-5 °C higher than that of the cold spot. The part bed mean temperature was lower than both spots at around 175°C, which is quite close to the set point.In terms of the cool-down, Figure 5a and 5b show plots of powder bed surface meantemperature change versus time via IR camera and processed by MATLAB.Figure 5: Polyamide 12 part bed surface temperature profile versus time during cool-down. (a) The mean temperature of part bed surface; (b) The part bed surface mean temperature profile processed by MATLAB.Figure 5a shows that the part bed surface temperature slowly decreased 5-10 °C in 350-380 seconds and then continually decreased to 120 °C in 800 seconds. Temperature at the moment that roller spread powder over the surface decreased by 15-30 °C.

Temperature at the moment that roller spread powder over the surface decreased by 15-30 °C. Figure 5b shows a smooth plot of the part bed surface temperature.To analyze nitrogen effects on the part bed profile, two experiments were conducted with the nitrogen purge gas on and purge gas off, respectively, which results are shown below in Figure 6.969Figure 6: Nitrogen circulation effect on part bed surface temperature. (a) The meantemperature of a part bed surface middle-line along the x-axis with purge nitrogen off at the moment nitrogen was off and 800 and 3600 seconds later; (b) The mean temperature of part bed surface middle line along the x-axis with purge nitrogen on, at nitrogen rates of 0.5, 1.5, 2.5 m3/h.Figure 6a shows that the part bed mean surface temperature increased with time after turning the circulation nitrogen off. The temperature range was selected from 165- 185 °C for better observation.

The temperature range was selected from 165- 185 °C for better observation. The whole part bed mean temeprature increased 2-3 °C after turning nitrogen off for 800 seconds and increased another 2-3 °C after 3600 seconds.Figure 6b shows the effect of different nitrogen circulation rates on the whole part bed mean surface temperature. The temperature increased 1-2 °C after changing the nitrogen rate from 0.5 m3/h to 1.5 m3/h, and remained when the nitrogen rate was changed from 1.5 m3/h to 2.5 m3/h. Data were selected 5 mins after changing the nitrogen rate. Within the first 5 minutes, nitrogen was continuously purging and heated by the heater. So the heated nitrogen had an effect on the part bed by increasing part bed mean surface temperature. However, the part bed mean surface temperature increased when the nitrogen rate stayed in a small range.

However, the part bed mean surface temperature increased when the nitrogen rate stayed in a small range. After changing the flow rate to 2.5 m3/h, the part bed temperature was the same as the part bed temperature when the flow rate was 1.5 m3/h.3. Laser pre-scan treatmentThe laser pre-scan experiment was conducted by building cubes on the part bed that were laser sintered by different powers ranging from 0.5 W- 7 W. The IR camera mounted on the SinterStation front window was used to record the temperature change. A thermograph during LS is shown in Figure 7.970Figure 7: thermograph from IR camera during LS. The array of 15 squares denotes LS build areas, each scanned at a different laser power setting from 0.5-7 watts.Cubes sintered by different power were highlighted by a different color, and each was separated by 0.25 inch in both x- and y-axes. Part bed operating temperature was 173 °C.

Part bed operating temperature was 173 °C. The detailed temperature rise and duration are shown in Figure 8.Figure 8: Temperature change of the sintered cubes by laser pre-scanning the powder bed surface. (a) Temperature rise as a function of LS power; (b) Pre-scanned part surface temperature as a function of power: before LS, after LS, and 10 seconds later after LS.971It is shown in Fig. 8a that the temperature rises roughly linearly and increases with power. Each datum was the temperature difference between the cube temperature before LS and the cube temperature immediately after LS. A linear relationship was empirically obtained for the temperature increase T:T = 0. 7958P (3)where T is the temperature difference from 0 to 6 °C, P is the laser power from 0 to 7 watts. The average residual is 0.53 in this case. It is reasonable since according to Eq.

7958P (3)where T is the temperature difference from 0 to 6 °C, P is the laser power from 0 to 7 watts. The average residual is 0.53 in this case. It is reasonable since according to Eq. (1), laser density energy increased linearly with power change, and temperature increment presents the laser density energy increment.Figure 8b shows the transient effect of the pre-scanned surface temperature change. It is seen that for each cube, the temperature increased 2-4 °C based on different laser powers. However, 10 seconds later, the pre-scanned surface temperature decreased back to the temperature before the scan, which means the effect dissipated quickly. This could be due to the low thermal conductivity of PA 12 (0.09-0.12 W/mK) [1] and convection.Figure 9 shows a thermograph of sintered tensile bars in the hot spot (right side) and cold spot(upper left side) during LS.Figure 9: Thermograph of sintered tensile bars in the powder bed hot spot and cold spot.

The temperature range was 165-190 °C.First, two tensile bars were sintered with 4 watts laser power in the hot spot and cold spot; after that, two more were sintered at the same position but with 10 watts laser power. Tensile stress and elongation was assessed using an INSTRON 3340 single column testing system. The original tensile bar gauge was 17 mm. Results are shown in Table 1.At the same position, the tensile bar sintered with larger power shows greater tensile stress. Moreover, tensile bars in the hot spot had larger tensile stress than that in the cold spot when the laser power was the same. The same situation happened in the case of elongation.

Moreover, tensile bars in the hot spot had larger tensile stress than that in the cold spot when the laser power was the same. The same situation happened in the case of elongation. These data demonstrate the importance of the powder bed surface temperature to achieve acceptable mechanical properties.972Tensile Stress (MPa) 39.3 20.4 54.3 41.6 Table 1: Mechanical properties of tensile bars sintered at the powder bed hot spot and cold spot with different laser power.Elongation (%) 21.7 6.5 27.6 24.7Hotspot_4W Coldspot_4W Hotspot_10W Coldspot_10WSummary and ConclusionVarious experiments were conducted to investigate and reduce thermal gradients on the part bed during LS. First, proper temperature set points for feed bins and part bed are important for LS operation. Second, thermal profile investigation provides the actual and MATLAB processed temperature profiles during the warm-up, LS, and cool-down stages.

Second, thermal profile investigation provides the actual and MATLAB processed temperature profiles during the warm-up, LS, and cool-down stages. Hot spot temperature was 3-5 °C higher than that of the cold spot in LS. Third, the part bed mean surface temperature increases with time after turning the circulation nitrogen off; it increased 1-2 °C after changing the nitrogen rate from 0.5 m3/h to 1.5 m3/h, and remained when the nitrogen rate was changed from 1.5 m3/h to 2.5 m3/h. Fourth, the laser pre-scan treatment shows the sintered part temperature increases linearly with increasing laser power, but the effect dissipates quickly after 10 seconds.

Fourth, the laser pre-scan treatment shows the sintered part temperature increases linearly with increasing laser power, but the effect dissipates quickly after 10 seconds. Last, mechanical properties were tested on tensile bars that were sintered in the hot spot and cold spot with different power, which demonstrates the importance of the powder bed surface temperature to achieve acceptable mechanical properties.AcknowledgmentsThe authors would like to acknowledge funding from the Air Force Research Laboratoryunder Grant #GRT00015778.References[ 1 ] M. Yuan, T. Diller, D. Bourell, 2011, ‘Thermal Conductivity Measurements ofPolyamide 12’, Solid Freeform Fabrication Symposium Proceedings (427-37), Austin, TX.[2]: J. Williams and C. Deckard, ‘Advances in modeling the effects of selected parameters onthe SLS process’, 1998, Rapid Prototyping Journal, Vol. 4, No.2, pp. 90-100.[3] S. Kolossov, E. Boillat, R. Glardon, P. Fischer, M.

ECTION FORCES AND FRICTION COEFFICIENTS FROM INJECTION MOLDING EXPERIMENTS USING RAPID TOOLED INSERTSM. E. Kinsella1, B. Lilly2, B. Carpenter2, K. Cooper3 1Air Force Research Laboratory, 2The Ohio State University, 3NASA Marshall Space Flight CenterAbstractExperiments have been performed with injection mold inserts made using solid freeformfabrication processes in an effort to further study such applications for economic production of small quantities of parts. Static friction coefficients were determined for HDPE and HIPS against P-20 steel, sintered LaserForm ST-100, and stereolithography SL 5170 using the ASTM D 1894 standard. Injection mold inserts were constructed of the same three materials and were used to inject cylindrical parts using HDPE and HIPS. Ejection forces were measured, and a model was used to calculate ejection forces and apparent coefficients of static friction.

Ejection forces were measured, and a model was used to calculate ejection forces and apparent coefficients of static friction. Statistical analyses were used to determine the effects of packing time, cooling time and packing pressure on ejection force for the three insert types. This paper compares experimental and calculated ejection forces, compares standard friction test results to calculated apparent coefficients of friction, summarizes the statistical results, and comments on the feasibility of using rapid tooled inserts for injection molding.BackgroundEquations developed for the ejection force for deep injection molded parts derive from , where FR is the ejection (or release) force, f is thethe friction-based concept coefficient of friction between the mold and the part, pA is the contact pressure of the part against the mold core, and A is the area of contact.

The version developed by Menges et al defines contact pressure asF Rf ×=p×AApA=)( TEd ×∆×rsmand therefore ejection force is:F R×=TEf )(d ×∆×rsm×L π2where E(T) is the modulus of the thermoplastic part material at ejection temperature, ∆dr is the relative change in diameter of the part immediately after ejection, sm is the thickness of the part, and L is the length of the part in contact with the mold core [1].Two important aspects of the above ejection force model are shrinkage and friction.Shrinkage results from thermal contraction and directional distortion [2]. Thermal contraction is669(1)(2)due to atomic vibration in which atoms move closer together at lower energy levels, and directional distortion results from orientation of polymer molecules during flow, and their subsequent relaxation back to a coiled state after flow ends.

Shrinkage is greatly influenced by ejection temperature and varies for amorphous and semicrystalline polymers. Semicrystalline materials exhibit greater shrinkage due to phase transformation of the crystalline portion. Amorphous polymers, on the other hand, contract much more gradually.Friction between the thermoplastic part and the injection mold core not only depends onthe mechanical relationship between the two surfaces, but also on an adhesive component inherent in the properties of the two materials at processing conditions. While the deformation (or mechanical) component of friction tends to be more easily defined, the adhesion component is rather more complex. In adhesive bonding, the surface tension of the adhesive is less than the free surface energy or critical surface tension of the adherend [3]. This allows the adhesive to wet and spread. Heat serves to increase the ability of the adhesive to absorb, dissolve, and disperse.

This allows the adhesive to wet and spread. Heat serves to increase the ability of the adhesive to absorb, dissolve, and disperse. Pressure and heat together improve wetting and spreading of more viscous materials. For polymers, the surface forces consist of van der Waals, coulombic and possibly hydrogen bonding forces [4]. The higher the surface free energy of the polymer, the greater the adhesive force. Very clean metal surfaces may promote chemical bonding.In this work, the ejection force, modulus, relative change in diameter, and thickness wereall determined experimentally as described below. Measured ejection forces were compared with those calculated from the model above, and standard friction testing provided a comparison for the calculated friction values.Testing and ExperimentationThe two rapid tooled injection mold inserts for this research were built using lasersintering and stereolithography processes.

The sintered insert was built using LaserForm ST- 100, a polymer-coated 420 stainless steel-based powder, which was sintered and subsequently infiltrated with bronze. The stereolithography insert was built using SL 5170 resin material from Vantico.The thermoplastic materials used for this research were high density polyethylene(HDPE), a crystalline polymer, and high impact polystyrene (HIPS), an amorphous polymer. The HDPE material was Lutene-H ME9180 from LG Chem. The HIPS material was BASF PS 495F. Both materials are commonly used for injection molding applications.Friction testing was conducted under three temperature conditions following the ASTM D 1894 standard, using the same thermoplastic and mold insert materials described above.

The tests were intended to compare coefficients of friction among injection mold insert materials, at elevated temperatures that more closely resembled processing conditions, and to friction coefficients calculated using experimental data and the ejection force model. The friction test matrix is shown in Table 1.670The first temperature condition was room temperature, 22 oC (71 oF). The second condition was ejection temperature, i.e., the temperature at which molded parts are ejected from the injection molding machine. The third condition consisted of first heating to an elevated temperature, then cooling down to ejection temperature before testing. While at the elevated temperature, a 0.9 kg (2-lb) mass was placed on the sled. When the plate cooled to ejection temperature, the weight was removed and the specimen tested. The purpose of the third temperature condition was to imprint the surface of the plate specimen onto the sled specimen.

The purpose of the third temperature condition was to imprint the surface of the plate specimen onto the sled specimen. This simulated the environment as well as the surface condition that occurs during ejection of a molded part.Elastic moduli for the HDPE and HIPS materials used in this research were measured atvarious temperatures using ASTM D 638 “Standard Test Method for Tensile Properties of Plastics.” Modulus results are shown in Figure 1.The experimental injection molded part was a closed-end, straight cylinder with a 32 mm (1.26 in) outside diameter, 49.6 mm (1.95 in) height, and 1.2 mm (0.05 in) wall thickness (Figure 2). The canister is designed with four vent holes in the base to prevent vacuum forces that would result during ejection from the core.A modular mold design was employed, consisting of a steel Master Unit Die (MUD)mold base and a core and cavity that can be removed and replaced with those of other materials.

Figure 3 shows the baseline core and cavity, made of P-20 steel, and the two rapid tooled core and cavity sets, made of SL 5170 epoxy resin and LaserForm ST-100 material. Each rapid tooled insert was machined to fit properly into the mold base.

This modification, however, did not completely eliminate the problem, and the experiments were completed using the P-20 cavity with the SL 5170 core. Process parameters for each insert and thermoplastic material combination are shown in Table 2.672Figure 2: Canister part with vent holes and no taper.Figure 3: Core and cavity inserts, before final machining, made of SL 5170 (left), P-20 steel (center), and LaserForm ST-100 (right).The mold was designed with one cavity and a heated sprue connecting directly to thebase of the canister. Core temperature was measured using three thermocouples positioned at different depths inside each core insert. Four subminiature load cells were used for measuring ejection force, one each positioned between an ejector pin and the mold ejector plate. A Sumitomo Injection Molding Machine, model SH50M, was used for the experiments.

A Sumitomo Injection Molding Machine, model SH50M, was used for the experiments. Digital imaging was used to measure inside and outside diameters of each part.A statistical experiment was designed to determine the effects of packing time, cooling time, and packing pressure on ejection force. In the experiments with P-20 and ST-100 inserts, there were two levels of packing time (2 and 6 seconds), three levels of cooling time (5, 10, and 15 seconds), and three levels of packing pressure (0, 5, and 10 percent of maximum, or 0, 10.93, and 21.87 MPa), eight repetitions each.

In the experiments with the SL 5170 insert, there were two levels each of packing time (2 and 6 seconds), cooling time (120 and 150 seconds), and packing pressure (0 and 5 percent of maximum, or 0 and 10.93 MPa), five repetitions each.673HDPE P-20 & ST-100 Inserts Temperature Profile:Velocity:35% or 56 mm/s (2.2 in/s)Sprue 210oCNozzle 210oCFront 199oCMiddle 193oCRear 177oCHIPS P-20 & ST-100 Inserts Temperature Profile:Velocity:40% or 64 mm/s (2.5 in/s) for P-20, 35% or 56 mm/s (2.2 in/s) for ST-100 Rear 191oCSprue 221oCNozzle 221oCFront 213oCMiddle 204oCHDPE SL 5170 Insert SL or P-20 CavityVelocity: Temperature Profile:25% or 40 mm/s (1.6 in/s)Sprue 177oCNozzle 177oCFront 171oCMiddle 166oCRear 160oCHIPS SL 5170 Insert SL or P-20 CavityVelocity: Temperature Profile:40% or 64 mm/s (2.5 in/s)Sprue 210oCNozzle 216oCFront 202oCMiddle 193oCRear 182oCTable 2: Injection molding parameters.Only a limited number of parts could be processed using the SL 5170 cavity due todeformation that caused sticking of parts.

At room temperature, the friction coefficient of HIPS was larger than that of HDPE on all three plate materials. Also, the friction coefficients of both thermoplastics on the SL 5170 plate were higher than those on the metal plates.Temperature did not make a dramatic difference in friction coefficient for HDPE on P-20,ST-100, or SL 5170. In all cases, elevating the initial temperature to imprint the surface of the specimen was presumed to increase the friction coefficient. It is noted, however, that the imprinted pattern was not very pronounced on the HDPE specimens. Therefore, the temperature to which the specimens were heated may not have been high enough to sufficiently soften the polymer.674The HIPS friction results were much more diverse than the HDPE results. HIPS on bothP-20 and ST-100 showed a slight decrease in coefficient from room temperature to ejection temperature, and an increase to elevated temperature.

HIPS on bothP-20 and ST-100 showed a slight decrease in coefficient from room temperature to ejection temperature, and an increase to elevated temperature. The increase was not significant for P-20, but was rather pronounced for ST-100.The HIPS friction coefficients increased with temperature on the SL 5170 plate. At ejection temperature, the friction coefficient is significantly higher than at room temperature, and at elevated temperature, the friction coefficient is dramatically higher. This extreme coefficient is attributable to adhesion, resulting from secondary forces between the two material surfaces [5].Injection Molding ResultsEjection force results from the injection molding experiments are shown in Table 4.Ejection forces for HDPE from the P-20 core, averaged per run, were generally lower than from the ST-100 core, which were lower than from the SL 5170 core.

Ejection forces for HDPE from the combination SL 5170/P-20 insert were higher than from the SL 5170 insert.In contrast to HDPE, ejection forces for HIPS from the P-20 core, averaged per run, are greater than from the ST-100 core for 12 out of 18 runs. Ejection forces from the SL 5170 core are much higher than from the other two. Also in contrast to HDPE, ejection forces for HIPS from the SL 5170 insert are larger than those from the combination SL 5170/P-20 insert.From the P-20 and ST-100 inserts, ejection forces for HIPS were higher than for HDPEfor all experimental runs. Two HDPE runs and three HIPS runs were completed with the SL 5170 core and cavity. For this insert, and for the SL 5170 insert with the P-20 cavity, ejection forces for HIPS were much higher than for HDPE.CalculationsEjection forces were calculated from the Menges model and compared with experimentalresults averaged across all runs (Figure 4).

Calculated values differ from actual values due in part to differences in measured and actual friction coefficients. Apparent coefficients of friction were calculated using the Menges model and compared with standard friction test results (Figures 5 and 6). Calculated values of the apparent coefficient of friction were higher than the standard test results due to differences in temperature and pressure.

Packing time had an effect on both thermoplastics with both rapid tooled inserts, and on HIPS across the board. Cooling time had an effect on both thermoplastics with the P-20 insert, and on HDPE only with the ST-100 insert. All three factors had an effect on HIPS with the P-20 mold insert. In the case of the baseline P- 20 insert, effects are very different between the two thermoplastics. For the two rapid tooled inserts, there is not as much difference in effects between the two thermoplastics.

Interaction effects are shown on the right side of the table.676Ejection Force HDPE400350300250N200150100500P-20ST-100SL/P-20SL 5170Experimental CalculatedEjection Force HIPSEjection Force HIPS800100007006008000500Experimental6000ExperimentalN400CalculatedNCalculated300400020020001000P-20ST-1000SL/P-20SL 5170Figure 4: Experimental compared with calculated ejection forces for HDPE (top) and HIPS.Static Friction Coefficient for HDPE1.601.401.201.000.80CalculatedStd Test0.600.400.200.00P-20ST-100SL/P-20SL 5170Figure 5: Coefficients of friction calculated from experimental results compared with those from standard friction tests for HDPE, all four inserts.677Static Friction Coefficient for HIPS6.005.004.003.00CalculatedStd Test2.001.000.00P-20ST-100SL/P-20SL 5170Figure 6: Coefficients of friction calculated from experimental results compared with those from standard friction tests for HIPS, all four inserts.HDPE Insert M aterial P-20 Steel SinteredPacking Tim e TpST-100 a P-20 a HIPS Insert M aterial P-20 Steel a Sintered ST-100 aSL 5170/Packing Tim e TpFactorsCooling Tim e Tc a aFactors Cooling Tim e Tc aPacking Pressure PpPacking Pressure Pp aTp-TcaTp-Tca aInteractionsTp-PpTc-PpInteractionsTp-PpTc-Ppaa aTp-Tc-Ppa aTp-Tc-Ppa aSL 5170/P-20 aaTable 5: Results from the designed experiment indicating which factors had a significant effect on ejection force.Other Observations of Rapid Tooled InsertsThe ST-100 insert held up well for the number of parts that were processed, and hadevery indication that it would last for many more.

There were differences in data between the ST-100 and P-20 experiments, however, indicating that there are thermal and adhesion differences that affect ejection force and friction coefficient.678The SL 5170 insert operated quite differently from the other two metal inserts. The resincore held up for more than 105 full parts, with minimal flashing and no catastrophic failure. A few anomalies with the SL 5170 insert were observed during processing, including adhesion, core swelling, internal defects, and deformation. Adhesion of the part to the SL 5170 core occurred with both thermoplastics, but to a much greater extent with HIPS.Defects internal to the SL 5170 core are shown in Figure 7. After the SL 5170 core processed about 40 parts, it began to show some internal defects, which looked like internal delaminations along the layers of the tool.

After the SL 5170 core processed about 40 parts, it began to show some internal defects, which looked like internal delaminations along the layers of the tool. The core, however, lasted through the entire experiment without these defects propagating into failure.Deflection of the SL 5170 cavity wall during injection had the greatest negative effect in the injection molding experiments. The wall of the cavity apparently deformed elastically under injection pressure. Then, although the part began to shrink, too much material had been forced into the mold, and was held fast by the cavity wall.In a preliminary study of this deformation, the SL 5170 insert was modeled using anABAQUS finite element analysis. Results show that the walls of the cavity do indeed elastically deform in a way that would cause parts to stick.

Results show that the walls of the cavity do indeed elastically deform in a way that would cause parts to stick. In Figure 8, deformation of the SL 5170 by HDPE injection, no packing, is concentrated near the base of the cavity, with a maximum magnitude of 0.06 millimeters (0.002 inches). Similar results were found with HIPS. When packing was included, the magnitude of the deformation of HIPS was greater and encompassed a greater area along the cavity wall.ConclusionsThe Menges model provides a reasonable method for estimating ejection forces and canpredict coefficients of static friction between the part and the core during ejection, given that accurate inputs to the model are available. Standard friction tests at elevated temperatures result in coefficient values that are closer to those during processing. The two environments, however, are not identical, so measured friction coefficients are not the same as those during processing.

The two environments, however, are not identical, so measured friction coefficients are not the same as those during processing. The coefficient of friction between HIPS and SL 5170 is very high due to excessive adhesion. Both the standard tests at elevated temperature and the experimental injection molding results show this.The ST-100 insert held up with no signs of wear or damage. For the given quantities, theST-100 performed just as well as the P-20. The SL 5170 insert is capable of limited processing of HDPE, but is not recommended for processing HIPS due to high adhesion. The core was durable enough to process more than 105 parts with no catastrophic failure. The major difficulty with the SL 5170 insert was deformation of the cavity during ejection, which caused parts to stick in the cavity when the mold opened.

The major difficulty with the SL 5170 insert was deformation of the cavity during ejection, which caused parts to stick in the cavity when the mold opened. Other issues with the SL 5170 core included core swelling and the development of internal defects after approximately 40 processing cycles.679Figure 7: Defects in the SL 5170 core.Figure 8: Simulation results of HDPE injection into SL 5170 insert, no packing.References1. Menges, G., Michaeli, W., Mohren, P. 2001, How to Make Injection Molds, Hanser Gardner Publications, Munich, Cincinnati.2. Burke, C., Malloy, R. 1991, “An Experimental Study of the Ejection Forces Encountered During Injection Molding,” ANTEC 1991, pp. 1781-1787.3. Cagle, C.V. (ed.) 1973, Handbook of Adhesive Bonding, McGraw-Hill, New York.4. Briscoe, B.J. 1981, “Wear of polymers: an essay on fundamental aspects,” Tribology International, August, pp. 231-243.5. Wu, S. 1982, Polymer Interface and Adhesion, Marcel Dekker, New York.680.

 EJECTION MECHANISM DESIGN METHOD FOR AIM TOOLSSundiata K. Jangha Graduate Research AssistantDavid W. Rosen Associate ProfessorThe George W. Woodruff School of Mechanical Engineering Georgia Institute of Technology Atlanta, GA 30332-0405404-894-9668david.rosen@me.gatech.eduKeywords: Injection Molding, Rapid Tooling, Ejection, Stereolithography, Decision-Based DesignABSTRACTOne of the key advantages of AIM tooling is time savings when producing small batch production quality parts. However, designing suitable ejection mechanisms is becoming a bottleneck. There are two goals of this paper. First, a model is presented that effectively characterizes the stresses on the mold core and part during injection molding. Second, a method is described for ejection system design. Our approach consists of a combination of analytical, computational, and physical experiments.

Second, a method is described for ejection system design. Our approach consists of a combination of analytical, computational, and physical experiments. The ejection system design method will first determine the feasibility of ejection for a particular part geometry, then will determine the number, sizes, and locations of ejector pins. Each phase of the method will be formulated into a Compromise Decision Support Problem, a multi-objective optimization problem formulation. An example will be presented to provide an idea of the robustness and the limitations of the method. Preliminary results indicate that this methodology is sound for a simple geometry.1.INTRODUCTIONRapid Erototyping (RP) is fast becoming the preferred method for hastening the product realization timeframe. The deficiency in this field is the lack of the ability to perform functional tests on the produced parts. Hence the development of the field.

The deficiency in this field is the lack of the ability to perform functional tests on the produced parts. Hence the development of the field. of Rapid looling (RT) which entails creating an injection molding tool capable of producing small batch parts, from an RP technology. RTcombinesthe speed of RP technologies with the realistic test specimens from injection molding. For this work We will consider AIM tooling which uses the Stereo;Lithography Apparatus (SLA) as the RP technology to create the molding tooL There are three phases to the injectiontnolding process; filling, cooling, and ejection. While all three of the phases present issues not considered when using traditional tooling, the final phase, ejection, has become the bottleneck for the RT process. As such, the ejection phase will be considered in this work.The practice of injection molding has always been based on heuristics. Hence the area needs some form of standardization.

As such, the ejection phase will be considered in this work.The practice of injection molding has always been based on heuristics. Hence the area needs some form of standardization. to ensure good ••parts are produced independent of the manufacturer. The most glaring area where thisis true is the· ejection of the parts. The material properties.of theSLAresin. are.much. closertolheinjectedluaterial.thanthose of t~e.••steeltools. Thus, the user must be more wary of the mold tool being damaged and even failing during the ejection of the part. We believe that the system that has been created will help to address these concerns.219The first step in the system is to determine whether or not the part can be ejected successfully. This step is carried out before any tools have been created. Therefore there is immediate timesaving. The second step in the system is to determine the actual ejection layout. This layout consists of the number, diameter, and locations of the ejector pins.

Therefore there is immediate timesaving. The second step in the system is to determine the actual ejection layout. This layout consists of the number, diameter, and locations of the ejector pins. This information will be invaluable to the user since there is less need for the incorporation of a highly skilled manufacturer in the early testing phase of product development.2. BACKGROUND2.1 Ejection Literature There has been quite a bit of research on ejection for traditional injection molds. This literature ranges from papers dealing with studies of the forces associated with part ejection [1-3], to sections from books describing the process in detail [4-7], to potential improvements to the ejection process [8]. However, until recently, very little work has been done with respect to RT molds. Clearly, the theory related to the forces associated·with ejection would apply. Yet, the rules developed would not due to the material property differences between RT and steel tools.

Clearly, the theory related to the forces associated·with ejection would apply. Yet, the rules developed would not due to the material property differences between RT and steel tools. The difference in material properties is. significant enough that it is necessary to redevelop ejection rules. Palmer [9], has established a foundation for these new rules. Her work entailed an investigation into the failure mechanisms for AIM tooling. During the course of this investigation she conducted extensive physical experiments the results of which were categorized into design rules. At the same time, CeDorge [10] has conducted research that investigates the thermal effects on the AIM molding process, and how that affects the life and productivity of the tool.The work that has been done in the area of ejection in rapid tooling has mainly been in the area of evolving design rules for traditional tools into usable rules for RT.

Dickens [11] has made strides towards this end by measuring the ejection force. characteristics for both pin and stripper plate ejection systems. At the same time, there are other areas that must be considered.2.2 Compromise Decision Support Problem The compromise DSP is a general solving. multi-objective, non-linear, framework for optimization problems [12]. Mathematically, the compromise DSP is a multi-objective decision model which is a hybrid formulation based on Mathematical Programming and Goal Programming [12]. The compromise DSP is used to determine the values of the design variables, which satisfy a set of constraints and bounds and achieve as closely as possible a set of conflicting goals. The structure of the cDSP is as follows:Given: Find: Satisfy: Minimize:A feasible alternative, assumptions, parameter values and goals. Values of design and deviation variables.

The structure of the cDSP is as follows:Given: Find: Satisfy: Minimize:A feasible alternative, assumptions, parameter values and goals. Values of design and deviation variables. System constraints, system goals, and bounds on variables Deviation variable that measures distance between goal targets and design pointThe alternative in this case is a part configuration. The goals of this system are to minimize the stress experienced by the part during ejection, and to minimize the number of pins to be used. The goals are often weighted depending on the intent of the designer. The design variables to be220investigated will be the number of pins, the pin diameters and the location of the pins. The orientation of the part and the location of the part with respect to the mold will also be considered but they are not design variables.

The orientation of the part and the location of the part with respect to the mold will also be considered but they are not design variables. The constraints to be satisfied in this system are the ejection rules discussed earlier..A solution to the compromise DSP is called a satisficing solution, because it is a feasible point that achieves the system goals to the "best" extent that is possible [13]. This notion of satisficing solutions is in philosophical harmony with the notion of developing a broad and robust set of top-level design specifications. The compromise DSP will be at the heart of the ejection system design methodology.2.3 Ejection Force Determination The ejection force (the force required to remove the part from the mold) is the primary consideration in the ejection process from a mechanics standpoint. The base equation for this force is as follows:FE =f.1*P*A Where Jl is the coefficient of friction, P is the contact pressure, and A is the contact area.

The base equation for this force is as follows:FE =f.1*P*A Where Jl is the coefficient of friction, P is the contact pressure, and A is the contact area. Glanvill [5] developed a more advanced formulation that considers an equivalent diameter:(1)FeLl *E*A*)]where:il<j> = "restrained" thermal contraction of plastic material across the equivalent diameter (m) <j>diameter of circle with a circumference equal to the length of perimeter of molding surrounding the male core (m)E = plastic modulus at the temperature of ejection (Pa)contact area (m2 coefficient of friction plastic coefficient of thermal expansion (m/m/°C))II a v = Poisson's Ratio Tm = "melt" temperature/softening point (OC) Te = part temperature at time of ejection (OC)This formulation is more robust than· the base equation, yet it still has limitations. The most glaring of these limitations are the lack of consideration of internal geometry and of draft.

The most glaring of these limitations are the lack of consideration of internal geometry and of draft. The ejection mechanism design method reported here is capable of dealing with both issues.3. EJECTION MECHANISM FORMULATION3.1 Approach There are two main areas that have to be dealt with in order to reduce the amount of time required to develop an ejection system. The first area to be understood is the nature of the force221(2)(3)that results from ejecting the part with respect to both the mold and the part. The second area to be understood is the effect that each ejector pin has on the entire system. This will allow for the addition or removal of pins as necessary. The overall approach to designing the ejection mechanism will be based on the Compromise Decision Support Problem (cDSP).

This will allow for the addition or removal of pins as necessary. The overall approach to designing the ejection mechanism will be based on the Compromise Decision Support Problem (cDSP). The word formulation for the cDSP is represented.· in Figure 1.GivenFindAssumptions -Injection Molding Parameters -Ejection Relations -Material PropertiesSy stem V ariables )Pin DiamiJt~r )NumOOrafPinsValues of Deviation VariablesSy stem Constraints>Pins must not. buckk >E1ns must notpillrcepcut >Maldccm, must !tOt/ail >RequirwJjorcff rnullt not ffxcud mackiM capaci{YThe design process for the ejection mechanism is depicted in Figure 2. The first step in designing the mechanism is to conduct a shrinkage analysis using an Finite Element Analysis (FEA). This analysis will dictate the amount of pressure that shrinkage due to the part's solidification will produce on the mold core. The second step is to formulate and solve the cDSP.

This analysis will dictate the amount of pressure that shrinkage due to the part's solidification will produce on the mold core. The second step is to formulate and solve the cDSP. The formulation of the cDSP requires only the addition of (part geometry, material properties). Once formulated the cD$P is solved using a computer algorithm. The output from this algorithm will be the number of ejector pins and their diameter. The next step will be to input the number of pins and their diameter into the algorithm that will be used to determine the potential ejector pin configurations. These configurations will then each be analyzed using an PEA.

These configurations will then each be analyzed using an PEA. From these configurations the final layout will be selected, completing the design of the ejector mechanism.System Goals)oMptjmlu Stre~pf'()ducedfHpart >Minimize numbtJrojpins -Bounds on System Variables>lI64"::::...Pin Dia""ulr~Part wail thickness )2 ~MmJbe,.ofpins .s /6Minimi~e -Deviation FunctionFigure 1 cDSP Word Formulationthe part-specificinformationthe In determination of part shrinkage, a two stage FEA is performed. The the first simulates solidification of the part via a thermal analysis. The second stage simulates the shrinkage of the part structural during cooling via analysis that uses the results from the analysis loading conditions.

The second stage simulates the shrinkage of the part structural during cooling via analysis that uses the results from the analysis loading conditions. After determining the the injection molding parameters, user the input required parameters along with the part geometry into the Finite Element Model this information, the PEA is run and the result pressure between the part and mold core.ordertofacilitatestageathermalasisto(PEM).WithisthecontactFinalResuitsFigure 2 Process Flowchart222Once the contact pressure is known, the cDSP can be solved. The cDSP has already been the completion of the formulated into a generalized form. formulation of the cDSP requires the user to input problem specific information (part geometry, part material properties, and injection molding parameters). The formulated cDSP is solved using a computer code written in FORTRAN. The algorithm conducts an exhaustive search of the design space.

The formulated cDSP is solved using a computer code written in FORTRAN. The algorithm conducts an exhaustive search of the design space. In this case the deviation function is composed of deviation variables representing the number of pins and the stress produced in the part. The more pins that are used the less the amount of localized stress. At the same time, the fewer the number of ejector pins required, the less expensive the process is with respect to both time and money. Hence the combination of design variables (number of pins, Pin Diameter) that produces the minimum value for the deviation function as a.whole is the desired satisficing solution. This solution is the first half of the ejector mechanism.

This solution is the first half of the ejector mechanism. The variable values from this solution are input into a separate algorithm to determine the other half of the ejector mechanism specifications, the pin locations.Like the shrinkage analysis,The algorithm that has been, created to determine the pin locations requires the information that is output from the solution of the cDSP. The user will have to specify certain information about the part. This includes the number offeatures that are to be ejected directly and the number of pins per feature. Based on the ejector pattern for the mold base, the algorithm considers a range of part orientations. and the resulting pin locations.• From this range the best set of pin locations is based on the desire to centrally locate the pins within the feature. Hence, for each angle, there is a part origin and a set of pin locations.

Hence, for each angle, there is a part origin and a set of pin locations. The output from this algorithm will be a list of all of the different part configurations (orientations and the, associated pin locations). Each one of these configurations will then be analyzed using an FEA ,to determine the configuration that produces the least amount of stress. This configuration will be ,the one selected. At this point, the specifications for the ejection mechanism are complete.vUJlA'::)"!.""!.,AA".::) take two forms. On one hand there are boundary conditions that must development of the ejection mechanism. On the other hand there are injection molding and part· design that must be taken into account. Both ofword formulation of the cDSP in Figure 1.this work, as is common with RTmolds, there is an ejector plate with is a hole at the center of the plate for the sprue. From the of intervals. One set is spaced at ~". The second set is intervals.

From the of intervals. One set is spaced at ~". The second set is intervals. As such, the mechanism design system must This compensation' takes place in the computer algorithm thatderived from well-known theory. First, the ejector to enforce this constraint is Euler's r·rI<.HTT'....223force can not exceed the capabilities of the machine. There is a set amount of force that the machine can generate in the assembly that ejects the part. The required ejection force can not exceed that amount of force. These constraints are addressed within the cDSP.In some of these cases, these constraints will never be violated. The concern. is for the cases where that is not true. Hence, the system has been designed such that the constraints are either met, or that particular combination of variable values is no longer considered.3.3 System Goals The two objectives for this system are to minimize the stress produced in the part due to ejection, and to minimize the number ofpins.

It is clear that these goals are in conflict with one another. However, that is what the cDSP was designed to handle.The minimization of stress within the part is crucial for "good" parts. There are a myriad of things that too much force could conceivably do to a part in an injection mold. Hence to avoid the part. any potential difficulty, the system has been designed to minimize the stress produced This is accomplished via the second PEA. The computer algorithm· dictates the potential configurations. Each· configuration is then analyzed· and the configuration that produces the lowest stress is the one that is selected. The natural reaction would be to add ejector pins so as to avoid stress concentrations. However, the addition of ejector pins is also an addition in both cost and time. Any increase in either one of these areas is undesirable. Hence the minimization of the number of ejector pins.4.

However, the addition of ejector pins is also an addition in both cost and time. Any increase in either one of these areas is undesirable. Hence the minimization of the number of ejector pins.4. RESULTS4.1 An Example Problem To demonstrate the implementation of the ejector mechanism design system, an example problem will be presented. Due to space limitations, all of the steps will not be described in detail. Thepart that will be used for this example is a box-shaped part (see Figure 3). The dimensions of the part depicted in Figure 3 are 5.08cm x 5.08cm x 1.27cm (2in x 2in x 0.5in).The first calculation step, denoted step 1 in Figure 2is the shrinkage analysis of the part and the mold. This two stage FEA is used to determine the amount of contact pressure between the part and mold. From the contact pressure, it can be deemed whether or not the mold core will fail when the part is ejected. The result of the shrinkage analysis is shown in Figure 4.

From the contact pressure, it can be deemed whether or not the mold core will fail when the part is ejected. The result of the shrinkage analysis is shown in Figure 4. As shown in figure, the part is shrinking on to the mold core magnitude of between 40 and 50 MPa. This is within the strength range of the mold material. This value, along with the experimentally derived value for the is incorporated into the coefficient of friction (0.29)Figure 4 Structural Contact (von Mises stress)algorithm that will be used to solve the cDSP.

The cDSP is to be solved as described in section 3.2.Part Orientation Pin DisplacementAngle450000 0.0000 45.0000OriainX 0.25 0 1.5y 0.75 0 0.5d11.011 1 ·0.7571d20.7571 1 1.111Pin1X 0.9649 1.0000 09646y 0.0351 0.0000 410354Pin Locations Pin3 Pin2X 2.1996 2.0000 3,6998Y -0.1289 1.0000 2.6998X 2.3791 1.0000 2.3789y 1.4493 2.0000 1.3789Pin4X 0.7854 0.0000 2.2856Y 1.2854 1.0000 1.2856Table 1 Abbreviated Outputfi'om Pin ConfigurationsThe formulated cDSP is solved by exhaustively searching the design space. The two system variables and the system variable bounds determine the design space as depicted in step 2 of Figure 2. The solution algorithm outputs a number ofpins\and the. pin diameters. Since<the example part has four features to be ejected and the wall thickness is 0.1 in, the results from the cDSP solution algorithm is four pins and a. pin diameter of 0.093875in. (3/32").

Since<the example part has four features to be ejected and the wall thickness is 0.1 in, the results from the cDSP solution algorithm is four pins and a. pin diameter of 0.093875in. (3/32"). This is to be expected due to the nature of the trade-Off In between number of pins and stress in the part. order to minimize the stress produced in the part, the ejector pin diameters should be maximized to minimize the local stress. One of the geometric constraints is that the pin diameters not exceed the wall thickness of the part. The va.lue output from the system is the largest standard pin diameter that will fit within the wall thickness. The next step is to determine the final configuration of the ejector mechanism. The output from step 3 of the flowchart in Figure 2 is a set of feasible pin configurations. This set ranges throughout the feasible range of part angles. with one feasible configuration emerging for each angle.

This set ranges throughout the feasible range of part angles. with one feasible configuration emerging for each angle. Table 1 is an abbreviated listing of the output.Figure 5 Resultant Stress Due to Pins Due to EjectionThese alternative configurations are run through an FEA to determine the configuration that yields the minimum stress in the part. This is solution step 4 in Figure 2. In this example the configuration that emerged was at 45 degrees. The stress in the part produced by this configuration is illustrated in Figure 5. The maximum stress in the part is between 50 and 58 kPa. This is well under the compressive strength of the material, which is 100 MPa. At this point the ejection mechanism is complete. The final layout is configuration 3 (45-degree orientation) in Table 1. The final layout with respect to the ejector plate is depicted in Figure 6.Figure 6 Ejector plate with final part configuration2255.

The final layout with respect to the ejector plate is depicted in Figure 6.Figure 6 Ejector plate with final part configuration2255. CONCLUSIONSAn ejection mechanism design method was developed that creates a.· mechanism while accounting for both the part geometry and the inJection molding parameters. involved with the product. The ••goalsofthe·.designsystem.f9rthismechanism are to.minimizethe.·numberof·pins, and the stress induced in the part due to ejeqtion. The intent of this syste11}j~tto allow the user to hasten/the realization ofaprod~5t.ibyshiftingthe production of prototype parts. that can be functionally tested to earlier inthedesign>prOP§ss.This is accomplished by the system in that it removes> a bottleneck from the.rapid tooling arena, and as such improves the ability of the user to ac~ievehig~(lu~lity\partsrelatively.quicklyand.with.aminimal amount of effort.

By using the methodS(tol.ltline.dinthispaper,theejection mechanism can be designed successfully .and with littlecosttothel.lserintiITle.This design system will help to keep the "rapid" in rapid tooling.it is notwithPlltlimitations..Therelitince on· the part having a. fairly simple geometry is •.afeatl.lre.Finally,as coIllplexityisaddedto the requirements, (increased. number of parts, number of features,etc ... )thesystembecomes computationally expensive. Effort~ are being made to furtherienhance the . capabilities and· eradicate the limitations of both the system and the underlying design methodology.ACKNOWLEDGEMENTSWe thankfully acknowledge the support of the Rapid Prototyping and Manufacturing Institute of the Georgia Institute of Technology and its member companies, and that of the George W. Woodruff School of Mechanical Engineering.REFERENCES1.2.3.4.5.6. 7.Malloy,R.,.P. Majeski. Design ofPin ejector Systems forlnjection Molds.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAN ELASTO/PLASTIC MODEL OF THE ULTRASONIC ADDITIVE MANUFACTURINGQ. Mao, N. Coutris, and G.M. FadelDepartment of Mechanical Engineering, Clemson University, SC 29630AbstractUltrasonic additive manufacturing (UAM) is a solid-state fabrication process that utilizes ultrasonic energy to bond metal foils layer by layer. It has been shown that bond formation is closely related to the plastic deformation of the material at the bonding interface. As a result, a material model is necessary to characterize the elastic and plastic deformation of material under UAM conditions.

As a result, a material model is necessary to characterize the elastic and plastic deformation of material under UAM conditions. The existing material models for UAM suffer from the following issues: 1) the models are established based on experimental data of a material different from the material used in UAM. 2) The models have not fully accounted for the influential factors in UAM. Therefore, a new material model is established by modifying the Johnson-Cook model which is known to be simple but effective to capture plasticity at elevated temperature and high strain rate. The model is based on experiments carried out on aluminum 6061-T6 which has been extensively used in the ultrasonic additive manufacturing. Ultrasound-induced effects including strain hardening, strain rate hardening, acoustic softening, and thermal softening are considered in the model.

Ultrasound-induced effects including strain hardening, strain rate hardening, acoustic softening, and thermal softening are considered in the model. In addition to plasticity, the bond formation/plastic deformation is also profoundly affected by the dynamics of the built feature which has not been considered in combination with plasticity in the existing studies. As a result, a UAM model is established in which both the mechanical and dynamic effects are taken into account. The model is shown to characterize well the plastic deformation at the bonding interface both qualitatively and quantitatively by comparing its predictions with experimental observations.IntroductionAlthough the UAM has only existed for a decade [1], its bonding process stems from the ultrasonic metal welding (UMW) which was invented over 50 years ago and is extensively implemented in industry. Numerous experimental investigations have been conducted to understand the bond mechanism of UMW [2-6].

Numerous experimental investigations have been conducted to understand the bond mechanism of UMW [2-6]. However, the mechanics behind the bonding process are still far from being completely understood. Particularly, the plastic shear deformation of materials at the bonding interface, which has been highlighted as a critical factor for bond formation [5], lacks full investigation and characterization. In addition, the dynamics of the built feature also have an influence on bond formation: the built feature undergoes resonance at a critical aspect ratio and reduces the differential motion between the top foil and the built feature. The loss of differential motion leads to loss of sufficient friction and thus results in bonding failure. Thus, the goals of this paper include: 1) characterizing the material behavior under UAM conditions and establishing a generic material model that is capable to predict the bonding mechanics.

Thus, the goals of this paper include: 1) characterizing the material behavior under UAM conditions and establishing a generic material model that is capable to predict the bonding mechanics. 2) Studying the influence of the change of macroscopic dynamics of the system on the bonding mechanics. 3) Relating bond mechanics to bond quality.317MotivationThe understanding of fundamental mechanisms of UAM bonding is limited. While it has been agreed by researchers that in general the bond formation undergoes three stages: 1) plastic deformation of asperities and break-up of surface oxides, 2) increasing contact area of pure metals with plastic flow enclosing the voids, and 3) bond formation and the maintaining of bonds through plastic deformation, yet the mechanisms behind bond formation is still arguable and even contradictory [2-5]. Early studies by Joshi showed that bonds are formed as the material undergoes diffusion, localized melting, and mechanical interlocking [7].

Early studies by Joshi showed that bonds are formed as the material undergoes diffusion, localized melting, and mechanical interlocking [7]. In support of Joshi, Gunduz et al. observed significant plastic flow at a strain rate of up to 103 /𝑠. The plastic flow on one hand initiates diffusion by increasing vacancy concentration, on the other induces localized melting through adiabatic heating [6]. Lee et al. observed mechanical interlocking between soft and hard material while bonding Cu to Ni-plated Cu [4]. Severe plastic deformation in copper is observed. According to Janaki Ram, however, the bonding process is universally due to solid-state metallic adhesion which involves adhesion of layers of atoms that create an apparent interfacial bond strength across the interface [3][8]. They further pointed out that the plastic deformation played a major role in all the three stages of bond formation as they repeat themselves in welding process.

They further pointed out that the plastic deformation played a major role in all the three stages of bond formation as they repeat themselves in welding process. Recrystallization is reported by several researchers [4] [9] [10]. Recrystallization is observed to be localized at the bond interface and is caused by heat generated from plastic deformation and interfacial friction. While the actual mechanism of UMW is arguable and could be material specific, it is evident that plastic deformation serves as a key component in bond formation. Without sufficient plastic deformation the bond would be degraded or even fail. If the plastic deformation at the bond interface can be quantified, it is possible to use it as an indicator for predicting bond quality and thus connecting UAM operating parameters (vibration amplitude, compression load, and weld speed) to bond quality through computation.

In order to quantify the plastic deformation, the bonding mechanics of UAM should be understood and a model be established.Pioneering work is done by de Vries who studied the mechanics of UMW, which shares a similar bonding principle with UAM but generates bonds only at contact points rather than along contact lines. They studied the internal stress within the top and bottom foil as well as frictional stress at the bond interface, and established a bond criterion that requires the internal stress to be greater than the yield strength of the material to initiate bonding. The internal stress is calculated based on both the yield stress which depends solely on temperature, and the compressional stress. Behind this mechanics model, several assumptions are made: 1) the plasticity is assumed to be perfect (free of hardening) and yield stress decreases as local temperature increases. 2) The friction coefficient at the bonding interface is assumed as a constant.

2) The friction coefficient at the bonding interface is assumed as a constant. 3) The ultrasonic sonotrode has a flat surface and thus the stress state is 2-D. While the mechanics model and the associated bond criterion provides insights to the understanding of UMW bonding, it is far from enough to characterize the mechanics in UAM. First, the yield stress under UAM condition is affected by multiple factors: plastic strain, high plastic strain rate, ultrasonic energy, and thermal energy. Second, the two mating foils undergo stick-slip friction which is affected by the dynamics of the built feature [11]. The friction coefficient is not constant but rather temperature-dependent [12]. Third, the stress state in UAM is not 2-D plane stress since the sonotrode is of cylindrical shape rather than flat.

The friction coefficient is not constant but rather temperature-dependent [12]. Third, the stress state in UAM is not 2-D plane stress since the sonotrode is of cylindrical shape rather than flat. Therefore the 3-D stress state has to be considered in the mechanics model.318Another factor that prevents the UMW model from being implemented for UAM is the difference in the dimensions of foils. While the dimension of the foils for UMW is flexible, the foils used for UAM process have fixed dimensions: 0.94 in. (23.88 mm) in width, 0.0059 in. (0.15 mm) in thickness, and customized length within the work space of UAM. Due to the small foil thickness, the dynamics of the top foil is generally insignificant when compared with that of the built feature and therefore the force required to excite the top foil is neglected.UAM models were also established to understand the UAM bonding mechanics [13-15]. In these models, the plasticity of the material models are of critical importance.

In these models, the plasticity of the material models are of critical importance. Siddiq and Ghassemieh modified the cyclic plasticity model proposed by Lemaitre and Chaboche to account for the acoustic and thermal softening effects [13] [16]. The cyclic plasticity model, however, is based on cyclic loading tests that are carried out at a strain rate (10−3/s) much lower than that of UAM (103~104/s), and on a material different from what has been used for UAM. As a result, the hardening effects due to strain rate are not taken into account. Additionally, the experimental data used for characterization of acoustic softening effect is from aluminum single crystal which is much softer than the aluminum alloy used in UAM. Kelly et al. utilized a power law function to characterize the stress-strain relationship in the plastic region for a UAM model bonding aluminum alloy 1100 [14].

Kelly et al. utilized a power law function to characterize the stress-strain relationship in the plastic region for a UAM model bonding aluminum alloy 1100 [14]. In the stress-strain relationship, the thermal softening effect is characterized by the exponential term and the acoustic softening by adding a factor to the power law function. Due to the lack of experimental investigation, the power law plasticity model is built on data of aluminum single crystal. The effect of strain rate is not considered. Siddiq and El Sayed modified a crystal plasticity model to account for acoustic and thermal softening in UAM [15]. The model parameters are optimized to fit the stress strain curves of aluminum single crystal. The model is then extended to polycrystalline aluminum based on EBSD studies of aluminum 6061-O. In this way, they try to understand the mechanism of acoustic and thermal softening in UAM at a microscopic scale.

In this way, they try to understand the mechanism of acoustic and thermal softening in UAM at a microscopic scale. Although the crystal plasticity model reflects the change in microstructure of the material, the number of model parameters significantly exceeds the number of observable parameters due to the complexity of the model and the lack of experimental investigation, and an optimization tool has to be introduced, which undermines the reliability of the model. In summary, the existing UAM plasticity models suffer from the following issues: 1) models are built on experimental data from a different test condition or a different material, 2) models are too simple to fully characterize or too complicated to realistically reflect the material behavior.

As a result, an appropriate material model should be established such that it can fully capture the material behavior by accounting for all the influential factors and its model parameters are sufficiently calibrated by experimental data of the correct material.While the plasticity governs the plastic stress in response to strain, the dynamics of the built feature also have a fundamental influence on the plastic stress through friction. As additional layers are deposited, the differential motion between the sonotrode and the built feature changes, which affects the interfacial frictional forces and thus the plastic shear deformation at the interface. Robinson et al. first documented the increasing deflection as the height of the built feature approaches its width (i.e. height-to-width ratio equals 1) and attributed the phenomenon to a lack of stiffness of the built feature with respect to the vibrating sonotrode [18].

height-to-width ratio equals 1) and attributed the phenomenon to a lack of stiffness of the built feature with respect to the vibrating sonotrode [18]. Alternatively, Zhang and Li associated the increase of deflection to a dynamic effect of the superposition of traveling waves imposed by the vibrating sonotrode [12]. Gibert et al. found that beyond the critical height- to-width ratio, bonding can be reinitiated and therefore the loss of stiffness is a dynamic effect319[11]. Furthermore, by studying the dynamics of the built feature they show that the built feature is close to resonance at the as height-to-width ratio approaches 1 under excitation of the ultrasonic sonotrode. As a result, the vibration of the built feature is synchronized with that of the sonotrode resulting in lack of differential motion for initiating frictional force and plastic shear force.

As a result, the vibration of the built feature is synchronized with that of the sonotrode resulting in lack of differential motion for initiating frictional force and plastic shear force. This change of dynamics has not been accounted in the studies of bonding mechanics and therefore its effect on bond formation is not clear.The Material ModelThe material under study is aluminum 6061, which is selected for its extensive implementation in UAM as well as the availability of data under a variety mechanical tests and studies. For the experimental investigation, aluminum 6061-T6 are purchased from Mcmaster Carr and machined to the shape of tensile specimens.

The total strain tensor 𝜀 can be decomposed into elastic and plastic strain tensors:𝑑𝜀 = 𝑑𝜀𝑒 + 𝑑𝜀𝑝In the elastic deformation region:𝜎 = 𝐶(𝜀 − 𝜀𝑝)The von Mises yield criterion and the associated flow rule are used to determine the conditions necessary for the onset of plastic deformation and the rate of change of plastic strain components as plastic deformation proceeds. The von Mises yield criterion requires the von Mises equivalent stress to be equal to the yield strength of the material for plastic deformation to take place. The yield function (criterion) and the associated flow rule are written:320(1)(2)f(σ̅) = |σ̅ − α| − (𝜎0 + 𝑅) = 0dεp = dλ∂f∂σ̅Where σ̅ is the von Mises equivalent stress, α is the back stress in kinematic hardening. Since kinematic hardening is not considered: α = 0. 𝜎0 is the initial yield stress, 𝑅 is the isotropic hardening term.

Since kinematic hardening is not considered: α = 0. 𝜎0 is the initial yield stress, 𝑅 is the isotropic hardening term. εp is the plastic strain of the material, 𝑑λ is a plastic multiplier.In this study, the Johnson-Cook (JC) model [19] is selected and modified to characterize the material plasticity for its simple mathematical form and extensive application. Although the model is completely empirical, it is effective in characterizing deformation of large plastic strain, in high plastic strain rate, and at elevated temperature. Additionally, the JC model is a particular type of von Mises plasticity model with an explicit form of hardening law and rate dependence. The original JC model is written [19]:𝜎 = (𝐴 + 𝐵𝜀̅𝑛)(1 + 𝐶𝑙𝑛𝜀̅̇ 𝜀̅̇ 0)(1 − 𝑇∗𝑚)̇0 is a reference Where 𝜀̅𝑛 is the equivalent plastic strain, 𝜀̅ ̇0 = 1 /𝑠. 𝐴 is the yield strength measured at zero-plastic- equivalent plastic strain rate defined as: 𝜀̅ strain, unit-plastic-strain-rate, and room-temperature.

𝐴 is the yield strength measured at zero-plastic- equivalent plastic strain rate defined as: 𝜀̅ strain, unit-plastic-strain-rate, and room-temperature. 𝐵 is the strain-hardening constant, 𝑛 the strain-hardening exponent, 𝐶 the strain-rate constant, 𝑚 the thermal-softening exponent, and 𝑇∗ a room-temperature (𝑇𝑟𝑜𝑜𝑚)-based dimensionless temperature defined as: 𝑇∗ = , where 𝑇𝑚𝑒𝑙𝑡 is the melting temperature. The first term defines the strain hardening, the second term defines the strain rate hardening, and the third term defines the softening at elevated temperature.

The first term defines the strain hardening, the second term defines the strain rate hardening, and the third term defines the softening at elevated temperature. ̇0, the JC model is reduced to: At room temperature (𝑇∗ = 0) and at the reference strain rate 𝜀̅̇ is the equivalent plastic strain rate, 𝜀̅𝑇−𝑇𝑟𝑜𝑜𝑚 𝑇𝑚𝑒𝑙𝑡−𝑇𝑟𝑜𝑜𝑚̇ = 𝜀̅𝜎 = 𝐴 + 𝐵𝜀̅𝑛A set of parameters for aluminum 6061-T6 can be obtained based on the quasi-static tensile test of aluminum 6061-T6: 𝐴 = 311.22 𝑀𝑃𝑎, 𝐵 = 240.11 𝑀𝑃𝑎, 𝑛 = 0.37.The JC model, despite its mathematical simplicity and modeling advantages, has a limitation to characterize plasticity with accuracy under the conditions in which high strain rate and ultrasound-induced softening are involved.

It has been shown by multiple investigations that metals with face-centered-cubic (f.c.c.) structure demonstrate a dramatic increase of dependence of dynamic flow stress on the instantaneous strain rate as the strain rate exceeds certain threshold (typically around 103 /𝑠) (fig. 1) [20][21]. It is believed that this abrupt change of strain rate sensitivity of the flow stress is due to the change of deformation mechanism. At low strain rate, the deformation is governed by the cutting or by-passing of obstacles by the dislocations. As the strain rate exceeds the threshold, the deformation starts to be controlled by phonon drag forces and the flow stress necessary to deform the material increases abruptly. This change of mechanism, however, is beyond the modeling capability of the standard JC model [20]. As a result, the JC model has to be modified such that both mechanisms are taken into account. Lesuer et al.

As a result, the JC model has to be modified such that both mechanisms are taken into account. Lesuer et al. proposed321(3)(4)(5)(6)a material model which utilizes different equations for characterizing the strain rate dependence based on the different mechanisms [20]. At low strain rate, a relation in form of Arrhenius equation is introduced based on the work of Frost and Ashby [23]:𝜀̇1 = 𝜀̇0exp [𝑄𝑘𝑇(1 −𝜎𝜏)](7)Where 𝜀̇1 represents the strain rate at which cutting and bypassing of obstacles by the dislocation is dominant, 𝜀̇0 is a reference strain rate, 𝑄 is the activation energy, 𝑘 is Boltzmann’s constant, 𝜎 is the flow stress, 𝜏 is the strength of obstacles at 0 K, and 𝑇 is temperature in Kelvin. At high strain rate, the relation is a power law equation:𝜀̇2 = 𝐶1𝜎𝐶2(8)Where 𝜀̇2 represents the strain rate at which the phonon drag force on dislocation is dominant, 𝐶1 and 𝐶2 are constants whose values are obtained through curve fitting.

The relations at low and high strain rates are combined to obtain the effective strain rate:𝜀̇𝑒𝑓𝑓 =𝜀̇1𝜀̇2 𝜀̇1 + 𝜀̇2(9)Similar material models can also be found in work by Sakino [21] and Manes et al.[24].The typical UAM processes involve strain rate at 103~104/s. The range of the shear strain rate can be calculated with the assumptions: 1) there exists no slip between the sonotrode and the top foil, i.e. the top surface of the top foil vibrates at the same amplitude as the sonotrode, 2) the bottom surface is fully welded to the built feature and is therefore fixed (fig. 2). The vibration amplitude 𝐴 varies between 5 to 50 µm while the thickness of the top foil 𝑡 is roughly 150 µm. As a result, the shear strain and strain rate of the top foil can be estimated:𝜏 =𝐴𝑡(10)𝜏̇ =𝜏𝑇/4(11)The calculated shear strain rate ranges from 2.66 × 103to 2.66 × 104 𝑠−1.

As a result, the shear strain and strain rate of the top foil can be estimated:𝜏 =𝐴𝑡(10)𝜏̇ =𝜏𝑇/4(11)The calculated shear strain rate ranges from 2.66 × 103to 2.66 × 104 𝑠−1. This range, however, falls around the turning point which indicates the transition of mechanisms in the rate sensitivity diagram. Since in real material, the transition of mechanism will not take place in all forest dislocations simultaneously, the sharp turn predicted by the proposed material model does not align with the experimental data.322Fig.1 The strain rate sensitivity diagram of aluminum 6061-T6. Experimental data are taken from [20] and [22].A = 5 ~ 50 µmSonotrodeP = 400N ~ 2kN µmTop foilt = ~150 µmBuilt featureFig.2 Schematics of the interaction between sonotrode and top foil.In order to better characterize the flow stress at transition phase of the mechanisms, more experimental data are studied to find a model that works better specifically for the UAM. Fig.

Fig. 3 shows the flow stress-strain rate diagram presented by Yadav et al. [25]. They collected experimental data of aluminum 6061-T6 from multiple studies and specifically examined the change of flow stress at the transition phase (fig. 3). Based on the data provided, a material model similar to that of Yadav et al. is implemented [25]:𝜎̅ 𝜎0̅̅̅̅= 𝐶 + 𝐷(𝜀̅̇𝜀0̅̅̅̇ )𝑏Where 𝜎0 is the stress at the reference strain rate 𝜀0̅̇ = 1 /𝑠. 𝐶, 𝐷, and 𝑚 are constants whose values are determeind by curve fitting: 𝐶 = 0.46, 𝐷 = 0.49, 𝑏 = 0.10. The comparison between the prediction of the model and the experimental data is shown in fig. 4. Based on the observation of little variation in shapes of the stress-strain curves at different strain rates, assumption has been made that the shapes of hardening curves at different strain rates are identical.323(12)Fig. 3 Strain rate sensitivity diagram of aluminum 6061-T6.Fig.

3 Strain rate sensitivity diagram of aluminum 6061-T6.Fig. 4 Comparison of strain rate sensitivity between experimental data and model prediction at strain rate range of the UAMIn addition to strain rate sensitivity, ultrasonic energy in UAM has been shown to have a significant effect on the plastic deformation of materials known as “acoustic softening”. Early documentation of this phenomenon is done by Blaha and Lengenecker, who observed instantaneous decreases and “zero plastic stress” in aluminum single crystals and a variety of metals as ultrasonic energy is applied to a specimen in tensile tests [17]. Moreover, it is shown from the stress-strain curves that the acoustic softening is similar to softening under heating. However, the energy consumption from the ultrasonic process is much less than that from the heating process. In order to better understand and characterize the softening effects, experimental investigations are conducted for aluminum 6061-T6.

In order to better understand and characterize the softening effects, experimental investigations are conducted for aluminum 6061-T6. The key observation is that the reduction of324the plastic stress is linearly proportional to the input of the ultrasound intensity (fig. 5). No macroscopic temperature increase is observed [26]. As a result, the JC model is modified such that the stress reduction due to ultrasound can be characterized as a linear function of the ultrasound intensity [26].𝜎̅ 𝜎0̅̅̅̅= 1 − 𝑑𝐼𝑢Where 𝜎̅ is the flow stress under ultrasonic irradiation, and 𝜎0̅̅̅ is the reference flow stress without ultrasonic irradiation. Both 𝜎̅ and 𝜎0̅̅̅ are taken as averages of the stresses measured at true plastic strain 𝜀𝑝 = 0.2 and 𝜀𝑝 = 0.3 in order to obtain an accurate measuring of stress reduction. 𝐼𝑢 is the ultrasound intensity, and 𝑑 is a constant that fits the model to experimental data.

𝐼𝑢 is the ultrasound intensity, and 𝑑 is a constant that fits the model to experimental data. At room temperature and reference strain rate, the value of 𝑑 is determined as 𝑑 = 1.2 × 10−4/J. Notice that in the experimental study of acoustic softening, the total ultrasonic energy 𝐸𝑢 can be directly measured by the ultrasonic welder power supply [26]. In the UAM process, while the sonotrode amplitude is known, a direct measure of ultrasonic energy is not applicable. In order to estimate the applied energy, a relation is introduced to associate the ultrasound intensity to the oscillation magnitude of the ultrasonic field. The relation is written as [27]:𝐼𝑢 = 𝜌𝐴̅2𝜔2𝑐Where 𝐴̅ is the equivalent oscillation amplitude of the ultrasound field. 𝜌 is the density of the medium, 𝑐 is the speed of sound in the medium, and 𝜔 is the angular frequency of the ultrasound, defined as 𝜔 = 2𝜋𝑓. 𝑓 is the frequency of ultrasound and is fixed at 20 kHz.

𝜌 is the density of the medium, 𝑐 is the speed of sound in the medium, and 𝜔 is the angular frequency of the ultrasound, defined as 𝜔 = 2𝜋𝑓. 𝑓 is the frequency of ultrasound and is fixed at 20 kHz. Since the bonding interface is very close to the sonotrode surface (roughly 0.1 mm), it is assumed that the oscillation amplitude of the ultrasonic field is the same as the amplitude of the sonotrode. Consequently, the ultrasound intensity can be characterized using vibration amplitude. By introducing equation (13) and (14), the acoustic softening can be incorporated into the stress-strain relation and thus the plasticity model (fig. 6). This characterization relies on the assumption that the amplitude of the ultrasound at the bonding interface equals the vibration amplitude of the sonotrode.

6). This characterization relies on the assumption that the amplitude of the ultrasound at the bonding interface equals the vibration amplitude of the sonotrode. In reality, however, intermittent slip takes place between the top foil and sonotrode, as well as between the top foil and the built feature which leads to an overestimation of the energy input at the bonding interface. On the other hand, the acoustic softening effect of the ultrasonic energy is underestimated in the experimental investigation due to the fact that a portion of the delivered ultrasonic energy is dissipated in the supporting structures rather than the in the specimen.325(13)(14)Fig. 5 The stress reduction is linearly proportional to ultrasonic energy input [26]Fig.

5 The stress reduction is linearly proportional to ultrasonic energy input [26]Fig. 6 The softening of aluminum 6061-T6 predicted by the model.The temperature effects on aluminum 6061-T6 is modeled using the original JC model based on the work of Banerjee and Bhawalkar [28] who studied plasticity of the aluminum 6061-T6 at the elevated temperature and strain rate. In the UAM process, the ambient temperature is typically controlled to 300˚F (around 150 ˚C). However, the local temperature at the bonding interface could be significantly higher due to interfacial friction and severe plastic deformation.

However, the local temperature at the bonding interface could be significantly higher due to interfacial friction and severe plastic deformation. Based on the experimental data provided, the exponential constant within the model is determined: 𝑐 = 0.68, and the melting temperature is set to 𝑇𝑚 = 1000 𝐾.326Fig.7 The comparison of experimental data and model prediction of the stress reduction due to temperature in aluminum 6061-T6.As a result, the modified JC model is obtained by incorporating the effects of high strain rate, acoustic softening, and temperature (equation (14)). The constants are given in table 2. Furthermore, the modified JC model and the classical JC model are evaluated in terms of their performances in characterizing different material behaviors under conditions of UAM process (Table 3).

In the next section, the modified JC model will be implemented to a finite element UAM model for understanding the material deformation in UAM bonding process.𝜎 = (1 − 𝑑𝐼𝑢)(𝐴 + 𝐵𝜀̅𝑎){𝐶 + 𝐷(𝜀̅̇ 𝜀̅̇ 0)𝑏}(1 − 𝑇∗𝑐)(15)𝑨/MPa𝑩/MPaTable 2 Constants for the modified JC model 𝒃𝑪𝑫𝒂𝒄𝒅/ J311.22240.110.460.490.280.100.681.2e-4Table 3 Comparison of classical and modified JC modelsStrain hardeningStrain rate hardeningClassical JC model Characterized based on quasi- static test Inaccurate prediction beyond strain rate of 𝟏𝟎𝟑 /𝒔Modified JC model Characterized based on quasi- static test Accurate prediction up to strain rate 𝟏𝟎𝟓 /𝒔Thermal softeningCharacterized based on tests at elevated temperaturesCharacterized based on tests at elevated temperatureAcoustic softeningNot availableCharacterized based on acoustic softening studies327Implementation of Material ModelThe modified JC model is implemented in the finite element commercial software ABAQUS 6.14 using the user-defined material subroutine.

The material model is integrated for computing the global stiffness matrix. Given the load increment, the stiffness matrix is then used for finding the displacement and strain increments that enable the system to reach quasi-static equilibrium. If equilibrium cannot be reached, the load increment will be reduced and the procedure is repeated. The steps are shown as follows. In the global level, the dynamics of the system is written as [28]:𝑴𝑢̈ + 𝑪𝑢̇ + 𝑲𝑢 = 𝒇𝑴 is the global mass matrix, 𝑪 is the global damping matrix, 𝑲 is the global stiffness matrix, and 𝒇 is external force vector. The dynamic equation is solved explicitly by decomposing each time step into a large number of small time increments. At each time increment, the quasi-static equilibrium is reached at each material point. The equation is written as:𝑲𝒖 = 𝒇𝑲 can be further decomposed:𝑲 = ∫ 𝑩𝑻 𝑫𝒆𝒑𝑩𝒅𝜴𝑩 is the deformation matrix, 𝑫𝒆𝒑 is the elastic-plastic constitutive matrix that is defined by the material model.

The equation is written as:𝑲𝒖 = 𝒇𝑲 can be further decomposed:𝑲 = ∫ 𝑩𝑻 𝑫𝒆𝒑𝑩𝒅𝜴𝑩 is the deformation matrix, 𝑫𝒆𝒑 is the elastic-plastic constitutive matrix that is defined by the material model. Write equation (17) into incremental form:𝑲𝜟𝒖 = 𝜟𝒇Since𝜟𝑢 = 𝑩−𝟏𝜟𝜺Where 𝜟𝜺 is the incremental strain. Inserting equations (18) and (20) into equation (19) yields:𝑩𝑇𝑫𝒆𝒑𝜟𝜺 = 𝛥𝒇Since 𝑫𝒆𝒑 depends implicitly on 𝜟𝜺, 𝜟𝜺 is found by integrating the constitutive law defined by the material model. Two integration schemes are available: explicit or implicit. In this work, both integration schemes have been explored and the explicit integration is implemented. The selection process is explained in the following section.

In this work, both integration schemes have been explored and the explicit integration is implemented. The selection process is explained in the following section. When 𝑫𝒆𝒑 is determined, the stress increment can be obtained accordingly using the following equation.328(16)(17)(18)(19)(20)(21)𝜟𝝈 = 𝑫𝒆𝒑𝜟𝜺 (22) Inserting equation (22) to (21) yields the incremental relation between 𝜟𝝈 and 𝜟𝒇 which leads to a differential relation. This relation can be integrated over the material domain and yields:∫ 𝑩𝑻𝝈 𝒅𝜴 = ∫ 𝒇 𝒅𝜴Let 𝒒 = ∫ 𝑩𝑇𝝈 𝑑Ω, then a force residual 𝒓 is defined:𝒓 = 𝒒 − 𝒇The force residual is used a convergence criterion for numerical integration and is checked at each iteration to see if equilibrium is achieved.To sum up, the general algorithm is as follow:1. Apply a load increment 𝛥𝒇, calculate for displacement increment 𝜟𝑢 and strain increment 𝜟𝜺.2. Compute the stress increment 𝜟𝝈3. Compute the residual 𝒓4.

Apply a load increment 𝛥𝒇, calculate for displacement increment 𝜟𝑢 and strain increment 𝜟𝜺.2. Compute the stress increment 𝜟𝝈3. Compute the residual 𝒓4. If the residual 𝒓 exceeds a set tolerance, set 𝛥𝒇 = −𝒓 and return to step 1.In step 1, 𝜟𝜺 can be obtained either through explicit or implicit scheme. The two integration schemes are explained briefly. For the explicit integration scheme, the strain difference is decomposed into elastic and plastic contributions:𝑑𝜺 = 𝑑𝜺𝒆 + 𝑑𝜺𝒑The relation between elastic stress and strain can thus be written:𝑑𝝈 = 𝑫(𝑑𝜺 − 𝑑𝜺𝒑)Where 𝑫 is the elastic stiffness matrix. First introduce the von Mises yield function for isotropic hardening:𝑓(𝝈, 𝜺𝒑) = 𝑓(𝝈) − 𝑟(𝜺𝒑) − 𝜎𝑦 = 𝟎Where 𝑟 is the isotropic hardening term, and 𝜎𝑦 is the initial yield stress.

𝐻 =𝑑𝒓 𝑑𝜺𝒑, 𝑟 is the isotropic hardeningterm. 𝑫 is the elastic stiffness matrix. 𝑓 is the yield function. 𝑑𝜆 is the magnitude of 𝑑𝜀𝑝. Apply forward Euler explicit integration scheme, equation (30) can be written in von Mises plasticity as:𝑑𝜆 =𝜎̅ − 𝜎𝑦 3𝐺 + 𝐻𝑑𝜀𝑒 (31)Where 𝜎̅ is the von Mises stress defined as 𝜎̅ = √3 2𝝈: 𝝈 and 𝜎𝑦 is the yield stress. The explicit scheme is relatively simple to implement and does not require the update of Jacobian matrix. However, it is conditionally stable and requires the time increment to be smaller than a critical threshold defined as:∆𝑡𝑐𝑟 ≈𝐿𝑐 𝑐𝑑(32)Where 𝐿𝑐 is the characteristic length of the element and 𝑐𝑑 is the dilatational wave speed.

With dΔ𝜆 the being computed,the stress and strain increments can be updated. The Jacobian matrix is then updated accordingly using the following equation:𝐉 =𝜕∆𝝈 𝜕∆𝜺(37)The implicit scheme is unconditionally stable since the yield criterion is imposed at each iteration. Therefore the time increment is not necessarily small and the increment length can be adjusted as iteration proceeds for optimum computational efficiency. However, the implicit iteration for computing Δ𝜆 and the update of Jacobian matrix 𝐉 are computationally costly.By implementing both the implicit and explicit integration scheme for solving the material model, it is found that the explicit scheme requires much less time than the implicit scheme for computing the stress-strain relation of a quasi-static tensile test. Moreover, the implicit scheme suffers from convergence difficulties as material and geometric nonlinearity increases.

Moreover, the implicit scheme suffers from convergence difficulties as material and geometric nonlinearity increases. The implicit scheme is more sensitive to mesh resolution, loading speed, and nonlinearity of the material model than the explicit scheme. As a result, the explicit scheme is selected for integrating the modified JC material model. A user-defined material subroutine (VUMAT) is developed for implementation of the model in ABAQUS.

As a result, the explicit scheme is selected for integrating the modified JC material model. A user-defined material subroutine (VUMAT) is developed for implementation of the model in ABAQUS. The work flowchart of the VUMAT is shown in fig.8.331ABAQUS Calls VUMATCalculate the plastic strain increment using the explicit equation: Eqn (31)Transfer state data from the last iterationUpdate the stressCalculate the elastic predictive stressCall VUHARD, calculate the yield stress with user- defined constitutive relationUpdate the state variables (strain, stain rate, yield stress), the internal energy, and inelastic energySatisfy von Mises yield criterion?YesNoUpdate stress with the elastic predictive stressEnd of subroutine VUMAT, return to ABAQUS main programFig.8 flowchart of VUMATValidation of Material ModelThe modified JC model is first implemented in simple finite element models for validation.

Specifically, three validations have been carried out: 1) 1-D tensile test subjected to ultrasonic vibrations with different amplitudes 2) 1-D tensile tests at high strain rates and 3) 3-D quasi-static tensile test. The predicted flow stress-plastic strain curves from the first two tests are compared with those from the analytical models (fig. 8). In the tensile test under ultrasonic vibration (fig. 8 (a)), the simulated curves run parallel to the analytical curves with slight deviation possibly due to error introduced by the explicit integration scheme. The maximum relative error for the prediction is 2.9%. In the strain rate test (fig. 8 (b)), the simulated stress-strain curves demonstrates serrated shapes at elevated strain rates possibly due to effects of dynamic loading as well as the elastic predictor-plastic corrector algorithm implemented in VUMAT. Despite the serrated shape, the simulated curves run parallel to the analytical curves and the relative errors are small.

Despite the serrated shape, the simulated curves run parallel to the analytical curves and the relative errors are small. The errors332are larger at elevated strain rates and the maximum value is 6.4%. Fig. 9 (a) shows the comparison between stress-strain curves from the 3-D quasi-static tensile simulation and from the experiment. The 3-D model is built such that the dimensions are the same as the standard tensile test specimen used in experimental studies (fig. 9 (b)). Only one quarter of the model is built for computational efficiency.

The model introduces a sonotrode, a top foil, and a built feature (fig. 10 (a)). Several simplifications are made in building the model. The sonotrode is modeled using rigid shell elements whereas the top foil and built features are modeled as deformable solid since the Titanium sonotrode is much harder than the foils. The laminar built feature is simplified as a parallelepiped assuming that no interfacial sliding takes place between two bonded layers. It should be noted that during welding severe plastic deformation is observed only in a region close to the weld interface. The region is characterized by the refinement and “flow” like morphology of the sub-grains and is typically 10 – 60 µm in depth from the weld interface [29]. As a result, two plastic deformation layers are introduced and attached to the mating surfaces of the top foil and the built feature respectively (fig.10 (b)).

As a result, two plastic deformation layers are introduced and attached to the mating surfaces of the top foil and the built feature respectively (fig.10 (b)). The elasticity and plasticity of the material in the plastic deformation layers differ from those of the bulk material of the top foil and the built feature. Specifically, acoustic softening only takes place in the layer and therefore the yield stresses are much lower in the layer than in the bulk material. The effective elastic modulus is also reduced by the surface asperities which have much lower stiffness than the bulk material. When under compression loading, the surface asperities are much easier to undergo plastic deformation than the bulk materials underneath. The effective modulus can be determined by assuming a Hertzian contact between two rough surfaces [30]:𝟏𝑬∗ =𝟐 𝟏−𝝂𝟏 𝑬𝟏+𝟐 𝟏−𝝂𝟐 𝑬𝟐Where 𝐸1, 𝐸2, 𝑣1, 𝑣2 are the elastic modulus and Poisson’s ratio of the two materials in contact.

For the contact between aluminum 6061-T6 foils, the effective modulus is roughly half of the elastic modulus: 𝐸∗ = 0.56𝐸. The lowered elastic modulus and yield stress allow for severe plastic deformation to take place at the weld interface. Both plastic layers are 20 µm thick according to observation made by Johnson et al. and the overall thickness of the top foil is 150 µm [29].334(38)SonotrodeTop foil Top foil plastic deformation layerBuilt feature plastic deformation layerTop foilBuilt featureBuilt feature(a) (b) Fig. 10 (a) UAM model overview (b) The plastic deformation layers between the top foil and built feature (the sonotrode is removed for clarity)Boundary conditions imposed on the sonotrode include a compression load of 1600 N in normal direction and a sinusoidal displacement of 20 µm at 20 kHz along the width of the foil horizontal direction. Clamping is imposed at the bottom of the built feature.

Clamping is imposed at the bottom of the built feature. As the sonotrode rolls along the length of the built feature, the part of top foil behind the rolling path is already “bonded” and thus are fully constrained to the built feature.Contact interactions are defined between sonotrode and top foil as well as between the top foil and the built feature. In the direction normal to contact interface, a surface-to-surface “hard” contact is defined which minimizes the penetration of the slave surface (soft material) into the master surface (hard material) at the contact locations using penalty method and impedes the transfer of tensile stress across the interface [28]. In tangential direction, a stick/slip criterion is imposed followed by a critical shear stress 𝜏𝑐𝑟 being defined such that the surfaces in contact stick as 𝜏 < 𝜏𝑐𝑟, and slip as 𝜏 > 𝜏𝑐𝑟. The value of 𝜏𝑐𝑟 is calculated based on isotropic Coulomb’s law: the product of friction coefficient and normal pressure.

The value of 𝜏𝑐𝑟 is calculated based on isotropic Coulomb’s law: the product of friction coefficient and normal pressure. The friction coefficient between the top foil and substrate is set to 0.4 based on previous experimental studies [12], whereas at the sonotrode-foil interface a “rough” surface contact is defined which imposes no slip at the interface. The assumptions of friction coefficient being a constant and the top being “grabbed” by the sonotrode without slipping are ideal assumptions for simplifying the problem. The influences of these factors will be covered in future work.In order to capture the severe deformation underneath the sonotrode with sufficient accuracy and at the same time reduce computational cost, a contact patch is generated with refined mesh for the top foil, built feature, and plastic deformation layers in between.

The 4-node linear tetrahedral element (C3D4) is selected due to its robustness over hexagonal shape and being able to adapt to relatively complex geometries. The linear elements is preferred over quadrature due to its robustness under contact conditions which leads to less convergence issues. The mesh size is determined based on the criterion of being able to predict modal frequencies with sufficient accuracy and capture the Hertzian contact stress distribution [33].The ABAQUS Explicit solver is used to interface with the user defined subroutine VUMAT for solving the model. The simulation is run on a high-performance computing (HPC) cluster with 16 processors for roughly 8 hours (Intel Xeon processors, 2.33 GHz). The simulation time is set to 0.3 ms which covers 6 vibration cycles of the sonotrode.

The simulation time is set to 0.3 ms which covers 6 vibration cycles of the sonotrode. Since the rolling speed (0.03 m/s) is much lower than the vibration speed (0.4-4 m/s), the sonotrode is assumed to be dwelling and only335ultrasonic vibration is considered in the simulation. Table 4 shows the operating parameters used for simulating the UAM process. The height of the built feature is varied such that the height-to- width ratio (H/W) is set to 0.2, 0.8, 1, 0.5, and 2 while the width and the length are fixed. As the height-to-width ratio varies, the dynamic interaction between the sonotrode and the built feature changes accordingly. Fig. 11 shows the sonotrode amplitude (dash line) dynamic response (colored solid lines) of the built feature at various H/W.

Fig. 11 shows the sonotrode amplitude (dash line) dynamic response (colored solid lines) of the built feature at various H/W. It is obvious that as the height-to-width ratio reaches 1, the response amplitude reaches the maximum in phase with the sonotrode, leading to insufficient differential motion between the built feature and the sonotrode for sustaining interfacial friction and forming bonds. At the height-to-width ratio of 0.2 and 1.5, the responses amplitudes are the minimum which is in agreement with the conclusion made by Gibert et al. that bonding can be reinitiated once the critical height-to-width ratio is exceeded [31]. At height-to- width ratio of 2, the response amplitude increases slightly possibly due to the loss of static stiffness as height increases.Table 4.

At height-to- width ratio of 2, the response amplitude increases slightly possibly due to the loss of static stiffness as height increases.Table 4. The operating parameters used for UAM simulation Vibration amplitude Vibration frequency Compression load Weld speed Temperature10 µm20 kHz1600 N0 m/s150 ºC1.50E-051.00E-05)m5.00E-06H/W=0.2( e d u t i l0.00E+00H/W=0.8H/W=1p m A5.00E-06H/W=1.5H/W=21.00E-05Sonotrode1.50E-050.E+001.E-042.E-043.E-044.E-04Time (s)Fig. 11 (a) UAM model overview (b) The plastic deformation layers between the top foil and built feature (the sonotrode is removed for clarity)The plastic deformation of both top foil and built feature are measured in the weld region in terms of von Mises stress. The measurement are taken from the plastic deformation layers of both the top foil and the built feature. The measuring points are located within the contact patch underneath the sonotrode (fig. 12). Fig.

The measuring points are located within the contact patch underneath the sonotrode (fig. 12). Fig. 13 shows the von Mises stress history of four different measuring locations at different height-to-width ratios. Several interesting observations can be made. First, at height-to-width ratio of 0.2, the von Mises stresses at all four measuring points exceed the yield stress which is indicated by the dash line. The actual yield stress (300 MPa) is lower than the nominal yield stress (350 MPa) under the effect of ultrasound and elevated temperature. At aspect ratio of 1.0, however, only one measuring point (edge of the top foil)336undergoes plastic deformation whereas three other points are all deformed elastically, which indicates insufficient plastic deformation at both measuring points. The results agrees well with the experimental observations made by Robinson et al [18].

The results agrees well with the experimental observations made by Robinson et al [18]. Second, by comparing the von Mises stresses at the edge and the center, it is shown that the stresses at the edge are generally higher than at the center in both top foil and the built feature. This could be due to the variation in differential motion across the width of the contact patch. The differential motion at the center are generally smaller than that in the edge thus causing the von Mises stress to be lower. This observation further points to the possible change of locations of bond degradation due to the change of height-to-width ratio.EdgeContact patchCenterFig.

(b) The von Mises stress history at H/W = 1.0337Conclusion and Future WorkIn this work, an elastic-plastic material model is developed and evaluated specifically for the UAM process. The model is derived by overcoming some of the limitations of the Johnson-Cook model when implemented for UAM. The material model is then implemented in a finite element UAM model for predicting material deformation at the bond interface under the influence of the dynamics and stick-slip friction of the system. The change of macroscopic dynamics of the built feature is shown to have a significant influence on the plastic deformation at the local bond interface. Specifically, the model first shows the significant change of dynamic response of the built feature at different aspect ratios. Then by comparing the evolution of von Mises stresses at different aspect ratios, the plastic deformation is shown to be significantly reduced at all measuring points at the critical aspect ratio (H/W = 1).

The stress evolution also shows a variation of plastic deformation across the width of the built feature. The material undergoes slightly more plastic deformation at the edge than at the center. These preliminary results have shown that the UAM model is capable of predicting material deformation at the bonding interface by considering a variety of influential factors: the dynamics of the system, strain and strain rate hardening, the acoustic and thermal softening, and stick-slip friction. The plastic deformation serves as an effective indicator of the bond quality and allows for the model to be able to quantify the bond quality based on the UAM operating parameters (amplitude, compression, weld speed).Following these preliminary results, the UAM model needs to be improved in several aspects: 1) a thermo-mechanical model is to be built to replace the current model in which the bonding process is assumed isothermal.

2) A temperature-dependent friction model needs to be established for more accurate prediction of thermal effects. 3) A set of criteria needs to be established to quantify the bond quality based on the prediction from the UAM model.AcknowledgementThis research was partially supported by the National Science Foundation under grant # 1068977 entitled “Understanding the Macroscopic Dynamics of Ultrasonic Consolidation”. The views presented in this work do not necessarily reflect those of our sponsor whose support is gratefully acknowledged. We also wish to thank Ms. Sophie Morneau from Branson Ultrasonics Corporation for loaning some ultrasonic components.Reference[1] White, D. (2003). Ultrasonic consolidation of aluminum tooling. Advanced Materials &Processes.[2] Kong, C. Y., Soar, R. C., & Dickens, P. M. (2003). Characterisation of aluminium alloy 6061 for the ultrasonic consolidation process. Materials Science and Engineering: A, 363(1-2), 99– 106.

In addition, PA 6 nanographene composites can also exploit the good mechanical properties of the parent polyamide 6, a structural thermoplastic ideal for 3D printing via fused filament fabrication (FFF). Hence, 3, 5, and 7 wt.% of NGP were blended with PA6 using co-rotating twin screw extruder to produce 1.75mm diameter for FFF using Lulzbot TAZ 6 3D printer. Scanning Electron Microscopy (SEM) was used to evaluate the degree of exfoliation of the nanographene and tensile and electrical resistivity test samples were manufactured via fused filament fabrication. The polyamide 6 nanographene composites were further subjected to annealing treatment at 80ºC, 140ºC, 200ºC and a comparison study was carried out on the observed tensile properties and electrical resistivity values of both annealed and not annealed treated.1.

Introduction1.1 MotivationThe scope of this research is to evaluate the effects on electrostatic charge dissipation at different annealing temperatures of nanographene platelets (NGP) added in polyamide 6 (PA6) filaments. Previous research of polymers filled with NGP suggests that the improved electrical conductivity properties are ideal for the parts manufactured by fused deposition modeling (FDM) also known as fused filament fabrication (FFF) that require electrostatic discharge protection. The size of the NGP platelets is 50 nm in diameter and the size of the filaments are 1.75mm in diameter. In this study, NGP were blended into PA6 using co-rotating twin screw extruder to compare electrical and mechanical properties and exfoliation of the nanoparticles. Scanning Electron Microscopy (SEM) was used to determine the degree of dispersion. Loadings of 3, 5, and 7wt% of NGP were evaluated in filaments that were not annealed and annealed at 80ºC, 140°C, and 200°C.

Loadings of 3, 5, and 7wt% of NGP were evaluated in filaments that were not annealed and annealed at 80ºC, 140°C, and 200°C. Tensile and electrical resistivity test coupons were prepared using FFF on a Lulzbot TAZ 6 3D printer and subsequently tested.10081.2Literature ReviewThere is a problem with the emission of electromagnetic waves that can interfere withsensitive electronic devices. Nonconductive plastics do not provide Electromagnetic Interference (EMI) shielding. A solution is to use conductive plastics where the material absorbs the electromagnetic waves by increasing the volume resistivity of the plastic. Conductive plastics are dependent on the material thickness, dispersion of the additive, and the conductivity of the additive. An advantage of using a conductive plastic instead of a coating is that the EMI shielding is integral to the part and cannot be removed by abrasion [1].Another problem with nonconductive polymers is electrostatic discharge (ESD).

ESD is the flow of an electrical charge that is stored in the nonconductive polymer to a material with a different electrical potential. Semiconductor and medical device industries need protection from ESD. Sensitive electronics manufacturers therefore establish electrostatic protective areas free of static electricity, using measures to prevent charging, such as avoiding highly charging materials by increasing the electrical conductivity of the polymer to that between an insulator and a conductor the material can provide ESD shielding [2].The packaging of these electrically sensitive devices can be fabricated with a conductive polymer to meet the ESD requirements. The ESD requirements of materials need to be antistatic (prevent a triboelectric charge), be able to dissipate an electrical charge with low surface or volume resistivity, or shield from electrostatic fields.

If the material is expected to have surfaces that will rub against each other it needs to have antistatic properties to not generate a static charge. If the material has a low enough electrical resistance it can dissipate the charge by grounding the electrical charge away from the sensitive electronics. To be categorized as a static dissipative the material needs a resistivity of greater than a conductor (1 x 104 Ohms) or less than an insulator (1 x 1011 Ohms). To have shielding characteristics the surface resistance needs to be less than 1 x 103 Ohms [3].FDM filaments with Polylactic Acid (PLA) polymer and reduced grapheme oxide (r-GO) have been used to make printed flexible circuits that can replace copper wire because of its high conductivity. 6wt% r-GO has shown a conductivity of the filament of 4.76 S/cm (0.21 Ohm*cm) [4].

6wt% r-GO has shown a conductivity of the filament of 4.76 S/cm (0.21 Ohm*cm) [4]. Black Magic is producing PLA with r-GO with a volume resistivity of 0.6 Ohm*cm to 3D print electrically conductive components [5].Various nanomaterials (e.g, Carbon Nanotubes (CNT), carbon black etc.) have been usedto make nanocomposites for the purpose of electrostatic charge dissipation. S. Athreya in 2009 made 4 wt% PA12/carbon black nanocomposites using Selective Laser Sintering (SLS) and showed that electrical conductivity and flexural modulus of the samples were higher than the samples made by injection molding [6]. Monica Calin in 2015 studied the electrical conductivity of PA6/MWCNT samples obtained by electrospinning and showed that volume resistivity decreased and surface resistivity remained unchanged with the applying load [7]. Joseph G.

Joseph G. Smith in 2004 prepared nanocomposites with space durable polyamides, SWCNT, and conductive additive (CuSO4 salt) that met the electrostatic charge dissipation requirements at space environment [8].1009Graphene Nanoplatelets (GNP) are flat nanosheets of carbon that is a high-quality cost-effective alternative to CNT. The lateral size is a few hundred nanometers to several micrometers and the thickness is 10 – 20 nm. With good dispersion the GNP will form a percolation network to form an ESD protection. Nylon is an engineered plastic and is a member of the polyamide group. It is a light weight, strong, flexible material and stable in chemicals, heat, UV light. The minimum quantity of filler required to form a conductive path within a polymer is defined as the Percolation Threshold (PT). A PT can be achieved with the addition of GNP of 5 wt% loading and above [9].In situ polymerization, solvent mixing, or melt mixing can be used to disperse GNP into PA11.

A PT can be achieved with the addition of GNP of 5 wt% loading and above [9].In situ polymerization, solvent mixing, or melt mixing can be used to disperse GNP into PA11. However, these dispersion methods have low yield and high production costs due to the energy required for dispersion and the large amounts of solvent that would be required in mass production. Melt Compounding in twin screw extruder is a dispersion method that is scalable, cost effective, fast, and would be a continuous mass production process. Filaments have been fabricated by Rashmi et. al. with PA11 with 0.5, 1, 3, 5 wt% GNP. They have reported filament volume resistivity of PA11 / 0.5 wt% GNP (4.54 x 107 Ohm * cm), PA11 / 1 wt% GNP (3.12 x 107 Ohm * cm), PA11 / 3 wt% GNP (2.27 x 107 Ohm * cm), and PA11 / 5 wt% GNP (1.92 x 107 Ohm * cm). If the properties of the filament can be transferred over to a FDM printed part ESD protection is possible [10].Polyamide 11 (PA11) / GNP have been studied by Tate, et al.

If the properties of the filament can be transferred over to a FDM printed part ESD protection is possible [10].Polyamide 11 (PA11) / GNP have been studied by Tate, et al. for ESD dissipation byusing the additive manufacturing method of SLS. GNP at 1, 3, and 5 wt% showed improvement in mechanical, thermal, and electrical properties. The GNP was dispersed using sonication in Isopropyl Alcohol (IPA). This sonication method produced a high degree of uniform dispersion and moderate exfoliation. Thin sheets of material were fabricated using compression molding. Scanning Electron Microscopy (SEM) was used to measure the quality of dispersion of the GNP. It was determined that less than 3 wt% GNP was needed to achieve low enough electrical resistivity for ESD applications. The PT for ESD applications is a surface resistance of 106 to 1010 Ohms/cm2 or an Electrical Resistivity of 104 to 108 Ohms*cm. PA 11 / 5 wt% GNP had an electrical resistivity of 2.46 x 106 Ohms*cm [11].Tate, et al.

PA 11 / 5 wt% GNP had an electrical resistivity of 2.46 x 106 Ohms*cm [11].Tate, et al. has researched the dispersion of 1, 3, 5, and 7 wt% NGP in PA11 by using aco-rotating twin-screw extrusion to be used in SLS. NGP has been shown increase electrical and thermal properties without degradation in the mechanical properties using SLS. The PT was increased to 7 wt% NGP to achieve an electrical resistivity on the order of 2.3 x 109 Ohms*cm with the injection molding process. The PT for volume resistivity is 1010 to 1011 Ohms*cm. The PT was increased to 5wt% NGP to achieve a volume resistivity 1.18 x 109 Ohm*cm with the compression molding process. It is believed that the twin screw extrusion process affected electrical resistivity from the high shear breaking down the NGP and decreasing the aspect ratio of the platelet.

It is believed that the twin screw extrusion process affected electrical resistivity from the high shear breaking down the NGP and decreasing the aspect ratio of the platelet. The sonication process had a better lower PT but will have limitations to scaling up for mass production [12].The drawback of the SLS process is that the machine cost is five times higher than a Fused Deposition Modeling (FDM) machine. PA11 / 3wt% NGP filaments were FDM 3D printed using a Lulzbot 4 3D printer. The electrical resistivity was measured using a Hioki Megaohmmeter at 500V with 30s of electric charge with a sample size 88 mm diameter 2.5 mm1010thick. The neat PA11 FDM printed polymer had a volume resistivity of 3.46 x 1013 Ohm*cm and an ultimate tensile strength of 4.59 MPa. The PA11 / 3wt% NGP that was FDM printed had a volume resistivity of 3.26 x 1013 Ohm*cm and an ultimate tensile strength of 11.73 MPa.

The PA11 / 3wt% NGP that was FDM printed had a volume resistivity of 3.26 x 1013 Ohm*cm and an ultimate tensile strength of 11.73 MPa. It was stated that the print quality was poor with interlaminar separation after printing. This study only looked at 3 wt% NGP which was not enough to reach the PT in the previously research using injection molding or compression molding. It was stated in the SAMPE conference proceeding that the electrical resistivity values of injection molded PA11 and 3wt% PA11/NGP nanocomposites were lower than 3D printed samples [13]. However if the data is reported correctly this statement is not true. The injection molded PA 11 / 3wt% NGP had a volume resistivity of 3.0 x 1014 Ohm*cm, the compression molded PA 11 / 3wt% NGP had a volume resistivity of 1.64 x 1014 Ohm*cm, and the FDM printed PA 11 / 3wt% NGP had a volume resistivity of 3.26 x 1013 Ohm*cm.

The FDM printed sample has ten times lower volume resistivity [12].In subsequent research the polymer was switched from PA11 to PA6 because of theavailability. PA6 / GNP were FDM printed using a Lulzbot TAZ 6 3D printer at 3, 5, 7 wt% GNP. SEM showed that the GNP structure was uniformly dispersed in the PA6. The tensile strength decreases with addition of GNP, tensile modulus increases at 3% wt but decreased at 5 and 7 wt%, and the % elongation increase with loading of GNP. All 3D printed parts showed interlaminar separation after printing. The volume resistivity of the FDM printed neat PA6 was 10.68 x 1013 Ohm*cm, PA6 / 3 wt% GNP was 11.62 x 1014 Ohm*cm, PA6 / 5 wt% GNP was 12.28 x 1014 Ohm*cm, and PA6 / 7 wt% GNP was 5.62 x 1014 Ohm*cm. The poor print quality created interlaminar separation which caused air pockets that increased electrical resistivity.

The poor print quality created interlaminar separation which caused air pockets that increased electrical resistivity. This study suggested using an IR light mounted ahead of the extruder for localized heating to fuse the layers together [14].1.3ApplicationsThe GNP has shown to have promise to make an ESD protected polymer material in PA6 andPA11 at a PT of 5 to 7 wt%. However, the FDM printing process of this material has produced parts that would not be acceptable for ESD protection. It was reported that the parts had poor interlaminar adhesion. Research needs to be conducted to improve the FDM printing process with these materials. It has been suggested that using Infrared Heating Lamps on the printed part could raise the temperature to eliminate the delaminations. If electrical resistivity can be decreased more, then we can use the nanocomposite for EMI shielding.

If electrical resistivity can be decreased more, then we can use the nanocomposite for EMI shielding. Since GNP is available at a broad range of platelet lengths, future research can be done to develop nanocomposites using GNP of various sizes.2. Experimentation2.1 Material SystemThe material systems used were polyamide 6 (PA6) in an extrusion grade homopolymer obtained from Advansix and nanographene platelets (NGP) M grade platelets obtained from XG sciences.10112.2 Manufacturing MethodsThe NGP was dispersed in the PA6 using melt compounding in a Werner Pfliederer co-rotating twin screw extruder. The screw diameter is 30 mm with a length to diameter (L/D) ration of 1:27. This dispersion method is scalable, cost effective, fast, and would be a continuous mass production process. 1.75 mm filaments were extruded for 3D printing using the fused filament fabrication (FFF) method. Filaments were made with 3, 5, and 7 weight percent NGP using a volumetric feeder.

1.75 mm filaments were extruded for 3D printing using the fused filament fabrication (FFF) method. Filaments were made with 3, 5, and 7 weight percent NGP using a volumetric feeder. The PA6 was dried in a desiccant drier before compounding. A Lulzbot Taz 6 was used to 3D print tensile samples that are shown in Figure 1.Figure 1. Lulzbot TAZ 6 Printer and Printed Tensile Samples2.3Characterization MethodsThe mechanical and electrical property characterization method of PA6/NGP was tensile(ASTM D638) and electrical resistance (ASTM D257). Samples were printed of the neat PA6 and 3, 5, and 7 wt% NGP. The tensile testing compared the tensile strength, tensile modulus, and percent elongation. Figure 2 shows the tensile testing setup and tested specimens. An extensometer was used to measure the strain in the tensile samples. Three specimens were testing for each material condition.1012Figure 2. Tensile Testing Setup and Tested Specimens.3.

An extensometer was used to measure the strain in the tensile samples. Three specimens were testing for each material condition.1012Figure 2. Tensile Testing Setup and Tested Specimens.3. Results3.1 Mechanical PropertiesTensile tests were performed on the neat, 3 wt.%, 5 wt.%, 7 wt.% NGP non-annealed samples, three samples were tested form each category. The annealing treatment (cid:68)(cid:87)(cid:3)(cid:27)(cid:19)(cid:4200)(cid:3)(cid:68)(cid:81)(cid:71)(cid:3)(cid:20)(cid:23)(cid:19)(cid:4200)(cid:38) does not lead to significant improvement in the tensile properties of the nanocomposites. However, it could be observed that annealing at 80(cid:4200)C exhibits significant increase in the percentage elongation at loading levels 3 and 5 wt%.

However, it could be observed that annealing at 80(cid:4200)C exhibits significant increase in the percentage elongation at loading levels 3 and 5 wt%. This is an indication of high ductility and could be attributed to the stress mitigation of highly heat affected areas, thereby distributing this out more evenly through the entire print, eliminating the possibility of fracture at specific region.Annealing at 200(cid:4200)C led to degeneration of the ultimate strength and elongation values and minor improvement in the modulus of elasticity. This could be as a result of the high temperature employed which is well above the glass transition of the parent PA 6 matrix. The glass transition of polyamide 6 ranges between 47- (cid:24)(cid:19)(cid:4200)(cid:38)(cid:3)and the 200(cid:4200)C annealing temperature is approaching the melting temperature of the PA6 which ranges between 223-250 (cid:4200)C.

There could have been certain thermal degradation occurring which in turn degrade the interfacial strength between the PA6 and NGP.Table 1 shows the summary of the tensile testing samples. The ultimate tensile strengthand modulus of elasticity was found to decrease with the addition of NGP. The percent elongation increased with the addition of NGP. The electrical resistivity decreased for the 80°C annealed specimens but increased for the 140°C and 200°C specimens. Figure 3 shows the ultimate tensile strength comparison. The highest ultimate tensile strength was found at the neat sample (33.44 MPa). Figure 4 shows the comparison of the modulus of elasticity. The modulus of elasticity decreased with the addition of NGP. Figure 5 shows the percent elongation comparison. The percent elongation increased with the addition of NGP. .1013Table 1.

Elastic Modulus for Annealing Temperatures and Wt.% NGP.1015% Elongation for Annealing Temperatures and % NGP12%10%8%C: .2~6%+-' !'CJ 01) C: 0 w 4%■ Not A nnealed■ 80C■ 140C■ 200C2%0%Neat3%5%7%% Nanographene Platelet (NGP)Figure 5. % Elongation for Annealing Temperatures and Wt.% NGP.The dispersion of the NGP in the PA6 and the morphologies of the not annealed andannealed at 80°C samples of the filaments were evaluated using scanning electron microscopy (SEM). The samples were sputter coated with three nanometers of gold. The filament samples were cut at a 45 degree angle before coating. The SEM images are taken on the cut surface. The representative micrographs are shown in Figures 6 to 11. All samples showed a good dispersion of the NGP in the PA6. The NGP are believed to be the rough looking areas and the smoother areas are the PA6.

All samples showed a good dispersion of the NGP in the PA6. The NGP are believed to be the rough looking areas and the smoother areas are the PA6. There was not a visible difference in the morphology between the annealed and not annealed samples.1016Figure 6. SEM – 3% Annealed at 80°C- -1µm - - Texas State Urnvers,Figure 7. SEM – 3% Not Annealed1017soonm Texas State Un,vers,tWO mag EB det 10/24/2018 0 34 nA 3 9 mm 30 007 x ETD 2 07 31 PMcurr- - -1µm - - - Texas State Un1vers1tFigure 8. SEM – 5% Annealed at 80°Ccurr WO mag EB det 11/29/2018 5 00 kV 86 A 5 0 mm 35 000 x TLD 1 26 55 PMHVFigure 9. SEM – 5% Not Annealed1018500nm Texas State Un1vers1t500nm- Texas State Un1vers1tcurr WD mag EE det 10/25/2018 5 00 kV 86 A 4 0 mm 30 020 x TLD 2 55 48 PMHV- - 1µm - - Texas State Un1vers1tcurr WD mag EE det 10/25/2018 - - 500 nm - -86 A 4 0 mm 50 030 x TLD 2 58 41 PMTexas State Un1vers1tFigure 10.

SEM – 7% Annealed at 80°C- 1µm - Texas State Un1vers1tWD mag EE det 11/29/2018 4 8 mm 34 968 x TLO 2 35 31 PMsoonm - Texas State Un1vers1tFigure 11. SEM – 7% Not Annealed4. ConclusionsTensile test specimens were 3D printed using the FFF method for the neat PA6, and with 3 wt%, 5wt%, and 7 wt% NGP and annealed at 80°C, 140°C, and 200°C. In conclusion the tensile strength and modulus of elasticity decreased with the addition of NGP in PA6 3D printed using the FFF method except for the 200° annealed specimens. The percent elongation decreased with the addition of NGP except for the 80°C Annealed increased up to 5% NGP. Addition of NGP in PA6 for FFF printing will increase the % elongation at the expense of decreased tensile strength and modulus. Electrical resistivity decreased for the 80°C annealed specimens to reach ESD percolation threshold. Electrical resistivity increased for the 140°C and 200°C annealed specimens.

Electrical resistivity decreased for the 80°C annealed specimens to reach ESD percolation threshold. Electrical resistivity increased for the 140°C and 200°C annealed specimens. SEM was used to verify the dispersion and morphology of filament samples that were annealed at 80°C and not annealed. The filaments showed good dispersion on all samples by using twin screw extrusion. It is unknown if annealing changed morphology as it was not a change observed in the SEM images. Some specimens were not flat. Heating profiles could be investigated to decrease warpage.10191.2.3.4.5.6.7.8.9.10.11.12.13.5. References(2009). Zeus Technical Newsletter EMI Shielding of Plastics. Zeus Industrial Products, Inc.: 8.Seal Science, Inc. Design Engineering & Precision Manufacturing - Brochure.Carter, S. Electrostatic Discharge (ESD) Properties of Plastic Packaging: Terminology, Standards, and Measurement.

‘As-built’ AM surfaces feature a high initial surface roughness (i.e. Ra > 10 µm), which often exceeds the specification for technical applications. In order to apply AM for highly stressed and cyclically loaded components, the as-built surface roughness needs to be reduced. Since conventional surface finishing processes as machining or blasting often show a limited applicability to complex shaped AM parts, an enhanced electrolytic polishing process was developed (3D SurFin®). Within the present study, Ti-6Al-4V AM plates and fatigue samples were produced in a powder bed laser beam system. The enhanced electrolytic polishing process led to a significant roughness decrease of approximately 84 % for a treatment time of 60 min. Also, a notable improvement of the fatigue performance of 174 % was achieved after a treatment time of 40 min in comparison to the as-built reference samples.1.

Also, a notable improvement of the fatigue performance of 174 % was achieved after a treatment time of 40 min in comparison to the as-built reference samples.1. IntroductionThe additive manufacturing (AM) technology is gaining increasing attention for actual industrial applications. Especially for the aerospace industry the processing of AM titanium components – and particularly the alloy Ti-6Al-4V (Ti64) – matters due to its beneficial corrosion resistance, its low density and the specific high strength [1]–[3]. Since the AM technology provides for a significant design freedom, topology optimized Ti64 components can be built by this layer-based manufacturing process leading to considerable savings of raw material, weight and therefore costs.

This has triggered various studies, that report about the mechanical properties [4]–[8], the microstructural development [9]–[13], as well as various applications [14]–[16] of additively manufactured components.Powder bed based laser beam melted (LBM) Ti64 parts are known to feature a high initial (‚as- built‘1) surface roughness of typically more than Ra = 10 µm right after the manufacturing process [8], [17]–[19], which strongly depends on the orientation of the surfaces in the build chamber. This effect can be a limitation for the application of cyclically loaded or highly stressed components, as well as for further post-processing procedures applied on the surface. Therefore,1 The ‚as-built‘ condition later on refers to the morphological aspect of the initial rough surface regardless of the thermal history of a specimen.2516the as-built surface roughness needs to be decreased.

Although different build-up strategies during the manufacturing process can be employed to reduce surface roughness [20], [21], additional surface finishing processes are typically needed [22], [23]. As presented elsewhere, the focus of the surface finishing process is still on conventional machining procedures, which are known to increase the fatigue performance in comparison to the as-built samples [24]–[27]. Since the applicability of machining is limited for complex shaped parts as is the case for most AM components, new surface finishing procedures need to be developed. Due to their wet-chemical nature, electrochemical processes are very attractive for smoothening of complex shaped parts [19], [26]. Conventional electropolishing processes have already been studied on AM parts previously and showed promising results with respect to roughness reduction [19], [22], [28].

Conventional electropolishing processes have already been studied on AM parts previously and showed promising results with respect to roughness reduction [19], [22], [28]. However, little is known about the enhancement of the fatigue performance of LBM Ti64 parts treated by these processes.Figure 1 shows a schematic of the experimental setup for an electropolishing process consisting of two cathodic counter electrodes, a direct current supply and the component as the anode. The electrolyte filling level, as well as the temperature can be controlled. A circulation pump is used to create a suitable convection within the electrolyte.+ -SampleElectrolyteFigure 1: Process flow scheme of the experimental test setup for electrolytic polishingFor each application, the electrolyte needs to be chemically adapted to the alloy to be treated. Upon applying a direct current, an ion movement is initiated in the bath.

Upon applying a direct current, an ion movement is initiated in the bath. During conventional electropolishing processes, roughness peaks are reduced when the flow conditions are such that transport through then near surface layer becomes the rate controlling step. However, the electrolytes used typically feature a high concentration of hazardous acids or alcohols [19], [22], [29], [30]. In addition, the treatment often comes along with substantial treatment times, or an undesirable high material removal leading to a considerable rounding of edges. Using alcohols in the electrolyte also induces the risk of explosion. Therefore, a new process based on the plasma electrolytic polishing process was developed, which will be referred to as the 3D SurFin® process in the following.

Therefore, a new process based on the plasma electrolytic polishing process was developed, which will be referred to as the 3D SurFin® process in the following. The main differences to conventional electropolishing are the higher applied voltage range of approximately 200 – 400 V, the shorter treatment times and the use of REACh compliant electrolytes that are essentially based on water as reported elsewhere [26], [31]–[34].2517Within this study, the influence of the treatment time of the 3D SurFin® process on the roughness of as-built LBM Ti64 plates was investigated. The focus was not only on the surface morphology but material removal and fatigue performance as well. Thus, fatigue samples were treated and compared with as-built samples to evaluate the effect of a smoothened surface on the fatigue performance.2.

Thus, fatigue samples were treated and compared with as-built samples to evaluate the effect of a smoothened surface on the fatigue performance.2. Experimental2.1 Sample manufacturingThe plates subsequently used for the characterization of the surface morphology were built at the Fraunhofer Institute ILT using a M270 laser beam system from EOS GmbH as described elsewhere [23]. All fatigue samples were manufactured by Liebherr Aerospace on a M290 laser beam system also from EOS GmbH. For all samples, TLS Technik GmbH supplied Ti-6Al-4V powder showing a spherical shape and an averaged particle size of 44 µm (Dp50). The nominal layer thickness during manufacturing was 30 µm. The plates were built with an energy density of 54.3 J/mm³, whereas an energy density of 55.6 J/mm³ was used for the build-up of the fatigue specimens. Prior to the removal from the build platform, all samples were stress relieved for 4 h at 540 °C under vacuum and cooled in an argon atmosphere.

Prior to the removal from the build platform, all samples were stress relieved for 4 h at 540 °C under vacuum and cooled in an argon atmosphere. Afterwards, all samples were hot isostatically pressed (HIP) for 2 h at 810 °C and 2000 bar in an inert gas atmosphere to reduce the risk of fatigue failure by inner pores or other imperfections.The characterization of the surface morphology prior and after the surface treatment was performed on plates with dimensions of 105 mm x 75 mm x 2 mm. They were built with an angle of 45° with respect to the build platform. According to [35]–[37], the surface facing the build platform at this angle shows the highest surface roughness without the use of support structures. Additionally, round fatigue samples were built vertically in z-direction under the same conditions in order to evaluate the effect of the surface treatment on the change in morphology and the corresponding fatigue behaviour.

As described in [23], a build-up angle different to 90° would result in a significant deviation from the sample geometry, as well as the symmetry axis affecting the load transfer during subsequent testing. According to DIN EN 6072 [38], the sample geometry FCE type 25A with a diameter of 5.64 mm was chosen (Kt = 1.035, length: 80 mm). Figure 2 shows the corresponding drawing of the net shape geometry. One series was built with this exact geometry, whereas two other series were built with a minor additional offset of 0.3 mm within the diameter to cover the material removal during the surface finishing process.

The threads (M16x1) were machined after the manufacturing, respectively after the surface finishing process.2518Figure 2: Round fatigue sample type FCE 25A (Kt = 1.035, length: 80 mm) according to DIN EN6072 [38]2.2 Surface finishingPrior to surface smoothening, the plates and fatigue samples were cleaned by an alkaline cleaning detergent (P3 Almeco from Henkel) for about 15 min at 65 °C ± 3 °C. Afterwards, all samples were immersion rinsed in deionized water for 3 min at room temperature and subsequently dried at 60°C ± 2 °C in a circulating air oven.The 3D SurFin® process was performed in a 200 l bath containing an electrolyte consisting of ammonium fluoride, sulphuric acid and a complex builder in accordance to [39]. The main ingredient of the electrolyte is deionized water with a mass percentage of > 80 wt.-%. As shown schematically in Figure 1, the electrolyte was continuously recycled by a pump and temperature was kept at 80 °C ± 3 °C.

As shown schematically in Figure 1, the electrolyte was continuously recycled by a pump and temperature was kept at 80 °C ± 3 °C. All samples were immersed vertically into the electrolyte with a set voltage of 300 – 350 V. Each of the three plates was treated sequentially to obtain a total treatment time of 2 min, 5 min, 10 min, 20 min, 40 min and 60 min. The two sets of fatigue specimens were separately smoothened for 20 min, respectively 40 min. After the treatment, all samples were immersion rinsed and dried according to the procedure after alkaline cleaning.2.3 Material characterizationFor the characterization of the surface condition prior to and after the treatment, the surface roughness was analysed according to DIN EN 4287 [40]. The roughness values Ra, Rz, Rt and Rv were measured using a Hommel Etamic Waveline profilometer T8000 with an accuracy of 0.002 µm.

The roughness values Ra, Rz, Rt and Rv were measured using a Hommel Etamic Waveline profilometer T8000 with an accuracy of 0.002 µm. All roughness values were determined three times in longitudinal direction on each side of the three plates.A scanning electron microscope (SEM) JSM 6320F from Jeol was employed to obtain high resolution micrographs. Each sample was sputtered with a platinum layer with a thickness of 4 nm prior to the SEM investigations. Additionally, an Auriga SEM from Zeiss equipped with a X-MaxN 50 detector from Oxford Instruments was used to determine the chemical composition of the surface by energy dispersive X-ray spectroscopy (EDX).

For titanium surfaces, the penetration depth was calculated to be 1.7 µm for the characterization of the chemical composition by EDX.2519The material removal was determined by differential weighting on a balance from Kern with a accuracy of ± 0.01 g.2.4 Fatigue testingThe fatigue samples were tested on two machines from Rumul. The samples with the as- built surface condition were tested on a Mikrotron test rig, whereas the electrolytically polished samples were tested on a Fractronic test rig. Both machines were equipped with a 24 kN load cell and calibrated prior to testing. The accuracy of the stress amplitudes were found to be within ± 1 % of the nominal values for both machines. The tests were performed at ambient temperature, a frequency of approximately 110 Hz and with an axial fatigue stress ratio of R = 0.1 at constant maximum stress levels.

The tests were performed at ambient temperature, a frequency of approximately 110 Hz and with an axial fatigue stress ratio of R = 0.1 at constant maximum stress levels. The fatigue limit was set to 3 x 107 cycles, which will be referred to as ‘run-out’ in the following.The fracture surfaces of broken fatigue specimens were analysed using the JSM 6320F SEM for a detailed investigation of fracture morphology and crack initiation sites.3. Results3.1 Surface morphologyAs described in [23] the as-built surfaces of additively manufactured Ti64 parts show a high initial surface roughness resulting from partially melted particles on top of the surface. This is demonstrated in Figure 3, which shows an irregular distribution of the particles leading to locally varying surface conditions. Moreover, it demonstrates a significant surface waviness, which can be classified as primary roughness caused by the melt pool during the manufacturing process [25].

Moreover, it demonstrates a significant surface waviness, which can be classified as primary roughness caused by the melt pool during the manufacturing process [25]. The surface shown in Figure 3 also features deep roughness furrows, which are hardly visible due to the particles on top of the surface contour.Figure 3: SEM micrograph of an as-built surface revealing partially melted particles responsiblefor the high initial surface roughnessThe averaged roughness values measured on the so called ‘downskin’ surface, which is facing the build plate, as well as on the ‘upskin’ surface on the opposite side are pointed out in Table 1 for the three plates. The Ra values show a difference of approximately 11 % between both sides,2520whereas the downskin face features higher roughness values in the as-built state.

The standard guidelines for most technical applications often call for much lower roughness values, and the results in Table 1 demonstrate that the as-built surfaces by far exceed the specifications.Table 1: Surface roughness values of the upskin and downskin side averaged out of three platesprior and after surface finishing (after 60 min)Ra, µmRz, µmRt, µmRv, µmUpskin as-built surface15.6 ± 1.3 118.9 ± 9.3 145.9 ± 20.0 51.2 ± 3.2Downskin as-built surface17.6 ± 0.7 131.5 ± 7.8 156.3 ± 12.6 56.3 ± 4.0Upskin surface after 3D SurFin®2.8 ± 0.5 19.4 ± 5.1 26.9 ± 9.18.9 ± 1.8Downskin surface after 3D SurFin®2.6 ± 0.3 17.0 ± 2.9 23.1 ± 6.18.0 ± 1.2Figure 4a) demonstrates the roughness decrease of the Ra value of the upskin, as well as of the downskin surface after different surface finishing treatment times. The averaged results for both sides of the plate out of all three plates were calculated.

The averaged results for both sides of the plate out of all three plates were calculated. The values after 60 min are also listed in Table 1 for both sides. Averaged over all three plates, the roughness could be reduced by about 84 % for a treatment time of 60 min. In addition, the initially different values of both sides became similar during the process. Although most of the change in surface roughness occurred in the first 40 min of the process, the material removal was still ongoing during the final 20 min (see Figure 4b)). Within that time, a weight loss of approximately 7.27 g ± 0.01 g resulted. In total, the plates lost an average weight of approximately 20.87 g ± 0.01 g.

Within that time, a weight loss of approximately 7.27 g ± 0.01 g resulted. In total, the plates lost an average weight of approximately 20.87 g ± 0.01 g. According to Figure 4b), the material removal follows a nearly linear trend over time despite the change in surface morphology.(a)(b)0510152025303540455055606502468101214161820Upskin surfaceDownskin surfaceRoughness Ra, µmTreatment time t, min05101520253035404550556065055606570758085Plate 1Plate 3Mass m, gTreatment time t, minPlate 2Figure 4: (a) averaged results for the roughness values Ra of the upskin and the downskin surfaces of all three plates after different treatment times for the 3D SurFin® process; (b) corresponding weight decrease2521Figure 5 shows the corresponding SEM micrographs after each treatment time from 2 min to 60 min. After 2 min, Figure 5a) already demonstrates an obvious attack on the single particles, as well as on additional roughness features.

After 2 min, Figure 5a) already demonstrates an obvious attack on the single particles, as well as on additional roughness features. Higher magnifications of this surface revealed an additional compact layer around the particles, as well as on top of the surface contour, which was partially dissolved. Applying a treatment time of 5 min (Figure 5b)), a significant roughness reduction was visible and most of the partially melted particles could be removed. To some extent the compact layer was still visible after 5 min, but fully vanished after a total treatment time of 10 min. As shown in Figure 5c), all particles could be removed after 10 min. and the microstructure of the substrate became visible. Although the SEM micrographs in Figure 5d)-f) indicate no further substantial change in morphological structure, isolated remaining roughness valleys as exemplary shown in Figure 5d).

Although the SEM micrographs in Figure 5d)-f) indicate no further substantial change in morphological structure, isolated remaining roughness valleys as exemplary shown in Figure 5d). Further material removal could decrease the furrows and led to a more homogeneous surface morphology (Figure 5f)).(a)(b)(c)2 min5 min10 min(d)(e)(f)20 min40 min60 minFigure 5: SEM micrographs of LBM Ti64 surfaces after different treatment times: (a) 2 min, (b)5 min, (c) 10 min, (d) 20 min, (e) 40 min, (f) 60 minA SEM micrograph of the LBM Ti64 surface after a treatment time of 2 min is shown at a higher magnification in Figure 6a). It details the initial powder particles, which were partially removed and coincidently encased with a material, which was different from bulk alloy. The thickness of this layer was approximately 3 - 6 µm. The corresponding EDX analyses are displayed in Figure 6b) and (c).

The thickness of this layer was approximately 3 - 6 µm. The corresponding EDX analyses are displayed in Figure 6b) and (c). As seen in Figure 6c), the central surface of the larger particle mainly consists of the elements expected for the bulk material Ti64 accompanied by some minor decontamination from the electrolyte (fluoride, sodium) and the environment (carbon). Figure 6b) reveals the element distribution of the layer formed on top of the particles. Besides typical contaminations (carbon, oxygen), this layer mainly consists of a titanium-fluoride compound and additional elements,2522which can be found in the electrolyte.

Besides typical contaminations (carbon, oxygen), this layer mainly consists of a titanium-fluoride compound and additional elements,2522which can be found in the electrolyte. Additional EDX analyses have been conducted in more areas on this sample, which all confirmed this trend.(a)(b)(c)Figure 6: (a) Backscatter SEM micrograph of the LBM Ti64 surface after a treatment time of 2 min; (b) EDX analysis of the surface layer formed on top of the roughness particles; (c) EDX analysis of a particle from the bulk material3.2 Fatigue resultsThe results of the fatigue specimens with as-built and after 3D SurFin® with 20 min and 40 min treatment time are displayed in Figure 7. The as-built test series demonstrated a substantially lower fatigue performance in comparison to samples with smoother surfaces. Two ‘run-outs’ were manually stopped after 3 x 107 cycles at 250 MPa and 350 MPa.

Two ‘run-outs’ were manually stopped after 3 x 107 cycles at 250 MPa and 350 MPa. A typical crack initiation site for this test series is shown in Figure 8 for a sample that failed after 4.1 x 104 cycles at a maximum stress level of 500 MPa.2523Fatigue results after 40 min treatment timeMaximum stress, MPaCylces to failureFatigue results after 20 min treatment timeFatigue results of samples with as built surfaces10410510610702004006008001000Figure 7: Comparison of the fatigue results of samples with as-built surface and with surfacesresulting from the 3D SurFin® process after treatment times of 20 min and 40 minIn Figure 8a) the crack initiation site is marked by a white rectangle. The higher magnification in Figure 8b) reveals a nucleation from a roughness valley at the outer surface contour.

The higher magnification in Figure 8b) reveals a nucleation from a roughness valley at the outer surface contour. This type of crack initiation was typical for all as-built samples.a)b)Figure 8: SEM micrograph of the fracture surfaces of a sample with an as-built surface condition that failed at a maximum stress of 500 MPa: (a) overview of the fractured surface, (b) detailed micrograph at the crack initiation siteIn order to investigate the influence of the surface finishing process on the fatigue behaviour, two additional test series with a treatment time of 20 min and 40 min were tested as seen in Figure 7. For both series, the fatigue life improved substantially compared to the as-built samples. A ‘run- out’ was obtained after 3 x 107 cycles at 400 MPa for the fatigue specimens smoothened after a treatment time of 20 min.

A ‘run- out’ was obtained after 3 x 107 cycles at 400 MPa for the fatigue specimens smoothened after a treatment time of 20 min. Figure 9a) demonstrates the most prevalent crack initiation site using the case of a sample broken after 1.6 x 105 cycles at a maximum stress level of 500 MPa. The2524nucleation site is again marked with a white rectangle. The higher magnification in Figure 9b) reveals nucleation from a roughness valley at the outer surface contour.The samples that were treated for 40 min using the 3D SurFin® process, achieved a more significant improvement of the fatigue performance compared to the treatment time of 20 min. In this case, three ‘run-outs’ were obtained at a maximum stress level of 700 MPa, 875 MPa and 890 MPa after 3 x 107 cycles.

In this case, three ‘run-outs’ were obtained at a maximum stress level of 700 MPa, 875 MPa and 890 MPa after 3 x 107 cycles. Taking the averaged maximum stress values of the ‘run-outs’ of the as-built and the 40 min polished test series into account, an increased fatigue performance of approximately 174 % can be calculated compared to the as-built condition.a)b)c)d)Figure 9: SEM micrograph of fracture surfaces of broken samples after the 3D SurFin® processwith a treatment time of 20 min ((a), (b)) and 40 min ((c), (d))Figure 9c) indicates the crack nucleation site of a sample treated for 40 min, which failed after 4.1 x 106 cycles at a stress level of 925 MPa. Again, the crack nucleation is marked by a white rectangle. As seen at the higher magnification in Figure 9d), the crack nucleated within the bulk material. This failure mode was prevalent for this entire test series.25254.

As seen at the higher magnification in Figure 9d), the crack nucleated within the bulk material. This failure mode was prevalent for this entire test series.25254. DiscussionAs reported elsewhere [8], [15], [19], the LBM Ti64 parts feature a high initial roughness resulting from the manufacturing process. The morphological difference between the upskin and the downskin surface of the plates built under an angle of 45° can be attributed to the so called staircase effect as described in [35], [36]. Consequently, as-built surfaces reveal significant roughness valleys and therefore notch effects, which cause a considerably lower fatigue performance as expected from previous reports [6], [17], [18]. The enhanced electrolytic polishing process (3D SurFin®) significantly reduced the surface roughness.

The enhanced electrolytic polishing process (3D SurFin®) significantly reduced the surface roughness. Although different roughness values Ra were initially measured for the upskin and the downskin side, the final values of both sides converged to an average roughness of Ra = 2.7 µm ± 0.5 µm after a treatment time of approximately 40 min. As demonstrated by the SEM micrographs, most of the features being responsible for the high roughness could be removed after approximately 10 – 20 min with the 3D SurFin® process. The material loss tended to show a linear decrease over time, i.e. the behaviour is similar to conventional electropolishing processes [19] in this respect.As presented in Figure 6a) - b), a 3 - 6 µm thick layer was built-up on top of the surface during the first 2 min. The layer mainly consists of a titanium-fluoride compound and a few additional elements, which are present in the electrolyte as well.

The layer mainly consists of a titanium-fluoride compound and a few additional elements, which are present in the electrolyte as well. This layer could also be found on the surfaces between and underneath the former powder particles. As it contains a considerable amount of titanium, the layer might result from the first dissolved material agglomerating or complexing close to surface areas with low electrolyte replacement due to less convection. The layer could be removed completely after a treatment time of approximately 10 min.Clearly, the as-built surfaces cause a substantial decrease in fatigue performance in comparison to conventional cast or wrought parts [2], [3]. By contrast, the smoothened fatigue samples demonstrated increased fatigue life, especially the specimens treated for 40 min.

By contrast, the smoothened fatigue samples demonstrated increased fatigue life, especially the specimens treated for 40 min. This can be attributed to the change in surface morphology as the SEM micrograph in Figure 5e) indicates that no roughness valleys and no initial powder particles remained at the surface after this treatment. This even led to an improvement of the performance compared to milled (hot isostatically pressed) LBM Ti64 samples [23], [25]. As indicated in Figure 9c), the broken specimens typically failed due to crack initiation in the bulk material but no longer by cracks formed at the surface.

As indicated in Figure 9c), the broken specimens typically failed due to crack initiation in the bulk material but no longer by cracks formed at the surface. The slight scatter within the results, which differ from the expected shape of the fatigue curve, might be referred to the fact that the applied stress amplitudes are close to the plastic deformation regime for hot isostatically pressed LBM Ti64 [7].By comparison, the roughness decrease after a treatment time of 20 min showed a substantially smaller fatigue improvement of approximately 33 %. From Figure 5d), it becomes apparent that particles were no longer present on the surface, but some considerable roughness valleys remained due to unsufficient material removal.

From Figure 5d), it becomes apparent that particles were no longer present on the surface, but some considerable roughness valleys remained due to unsufficient material removal. Therefore, a treatment time of at least 40 min is needed in this process to obtain a balanced relation between material removal and increase in fatigue performance.This effect leads to the assumption that there is no direct correlation between the fatigue performance and the individual roughness values. This effect was previously shown in [23]. The roughness valleys in Figure 5d) might be missed by tactile roughness measurements leading to2526more optimistic results. Alternative surface characterization methods such as imaging techniques or dye penetrant inspection could be used to assure the absence of surface defects.5.

Alternative surface characterization methods such as imaging techniques or dye penetrant inspection could be used to assure the absence of surface defects.5. ConclusionAdditively manufactured metallic components feature a high initial surface roughness right after the manufacturing process, which needs to be decreased for further post-processing as well as for the application in cyclically loaded components. In the present study an enhanced electrolytic polishing process (3D SurFin®) was applied on LBM Ti64 plates and fatigue specimens in order to evaluate its efficiency in terms of roughness reduction, as well as improvement of fatigue performance. Based on the results, the following main conclusions can be drawn:A high initial surface roughness with an averaged Ra of 16.6 µm ± 1.0 µm was measured on the LBM Ti64 plates for the as-built surface condition.

This is a result of partially melted powder particles, as well as a high surface waviness caused by the manufacturing process.The notch sensitive LBM Ti64 samples with the as-built surface condition displayed a deteriorated fatigue performance with two run-outs at 250 MPa and 350 MPa after 3 x 107 cycles.The application of the 3D SurFin® process could reduce the initial roughness to a final averaged value of Ra = 2.7 µm ± 0.4 µm within a treatment time of 60 min.Treating fatigue samples by the enhanced electrolytic polishing process, the fatigue performance could considerably be increased.

For a treatment time of 40 min, the technical endurance limit was around 820 MPa after 3 x 107 cycles, which is similar to the value obtained for hot isostatically pressed, milled LBM Ti64 fatigue samples.A treatment time of 20 min resulted in a lower improvement which could be attributed to an insufficient material removal leaving occasional roughness valleys, which could still cause a crack initiation from the outer surface contour.AcknowledgementsThe present study was supported by the German aeronautic research program (LuFo) within the framework of the project GenFLY. The authors thank the project sponsor Airbus, as well as the corresponding project partners Liebherr Aerospace and Fraunhofer ILT. The corresponding author also thanks Norbert Schupp, Max Kolb, Vitus Holzinger, Christian Plander and Freerk Syassen for their support.References[1] [2][3][4]Lütjering G, Williams JC. Titanium. 2nd ed. Germany: Springer; 2007. Campbell FC.

ectrochemical Infiltration of Laser Sintered Preforms with MetalsAbhishek Goel and David Bourell Laboratory for Freeform Fabrication, The University of Texas at Austin 1 University Station, MC C2200 Austin, Texas 78712 USA +512-471-3170 voice, +512-471-7681 fax agoel@mail.utexas.edu, dbourell@mail.utexas.eduAbstractThe proposed work deals with room temperature infiltration of selective laser sintered preforms with metals. The approach adopted for carrying out room temperature infiltration was adaptation and modification of electrochemical deposition techniques. Electroless deposition and Electrolytic deposition processes were modified in order to carry out metal ion infiltration and their deposition on the walls of the interconnected pores. The electrolytic infiltration process was modified by inserting a conductive graphite cathode in the center in order to attract positive nickel ions through the interconnected porous network and hence deposit them on the walls of the pores.

One of the major advantages of electrochemical infiltration is the lower processing temperature. Low temperature reduces both energy consumption and associated carbon-footprint and also minimizes undesirable structural changes. Both conductive and non-conductive preforms can be electrochemically infiltrated and MMCs produced by this method have potential for use in structural applications. This research was sponsored by the National Science Foundation, Grant DMI-0522176.IntroductionThis work deals with electrochemical infiltration of parts fabricated using one of the most common powder processing additive processes, Selective Laser Sintering (SLS). Powder metallurgy and melt infiltration are the most common techniques for producing metallic matrix composites (MMCs). A common feature of these techniques is high processing temperature which is a major cause for high production costs, poor mechanical properties and high sustainability penalty.

A common feature of these techniques is high processing temperature which is a major cause for high production costs, poor mechanical properties and high sustainability penalty. Electrochemical deposition is a technique which is most commonly used to deposit a thin film coating on various types of material substrates. In the following work the principles of electrochemical deposition are used to carry out electrochemical infiltration of metal ions into sintered preforms. One of the major advantages of electrochemical infiltration is the lower processing temperature, normally room temperature, which helps to avoid undesirable chemical and structural changes. When electrochemical techniques are applied to form MMCs the working conditions become much more complicated where the electrolyte must transport to the interior porous structure through narrow channels and at the same time the by-products of reaction must be driven out of the preform.

The work aims at facile, low cost room temperature infiltration of selective laser sintered preforms to create fully functional dense parts. Selectivelaser sintering (SLS) is a commercial, powder-based manufacturing process that produces parts with arbitrary shape based on a computer solid model. Figure 1 shows a schematic of the process. Powder is spread as a thin layer approximately 125 microns thick. A computer- controlled laser scans the surface and melts together the powder, creating a thin cross section of the part. Alternatively, the powder may be mixed with a transient binder which melts under the scanning laser beam and is removed or converted in a post-processing step. This process is repeated thousands of times to create an entire part with virtually unlimited geometric complexity.Figure 1. Schematic of Selective Laser SinteringSLSed parts are typically 50-65 percent dense with virtually all porosity being open or connected.

Schematic of Selective Laser SinteringSLSed parts are typically 50-65 percent dense with virtually all porosity being open or connected. The average pore size is typically the order of the powder size, approximately 30-60 microns. In electrochemical deposition, this pore network is filled with an electrolytic liquid, and deposition of a solid phase is initiated internally. The time required to close a pore is short since only a pore radius (~15-30 microns) needs to be deposited for closure. By control of the deposition process spatially, the electrolytic liquid can be successfully ejected in front of the deposition. The general steps for electrochemical infiltration are similar to traditional SLS infiltration. First, the non-metallic/metallic phase is mixed in powder form with an appropriate binder powder. The mixture is then SLSed to create a porous “green” part bound by the binder. After this the green part is converted to a “brown” part by heating to dissociate the binder.

The mixture is then SLSed to create a porous “green” part bound by the binder. After this the green part is converted to a “brown” part by heating to dissociate the binder. In this way organic ligaments or bridges between individual powder particles are replaced by a residue of equal or greater bond strength. The brown part is then infiltrated by dipping it in a bath containing a metallic salt as an electrolyte. Metal ions diffuse into the porous network of the brown preform where they combine with a counter electron flow resulting in metallic deposition within the pore network [1-4]. Silicon carbide is an engineering ceramic for which the indirect SLS process has a great potential. This material is of interest where high temperature strength and stability are required, and where wear and corrosion are issues. Some of its numerous applications include electronics processing wafer carriers and boats, kiln components and wear parts.

Some of its numerous applications include electronics processing wafer carriers and boats, kiln components and wear parts. However, silicon carbide objects, and non-metallics in general, are often difficult to manufacture due to their high melting temperatures (around 2700°C for SiC), poor sinterability, limited ductility and difficulty to machine. Selective laser sintering coupled with electrochemical infiltration offers a cost-effective and rapid manufacturing mechanism for metal-SiC composites. SiC is one of the potential candidates for high temperature structural components. However, the most crucial aspect that limits the application of SiC-based ceramics is their brittleness. SiC isalso quite difficult to densify because of the covalent nature of the Si-C bonding and a very low self-diffusion coefficient [5, 6]. Even liquid phase sintering is not able to provide highly dense SiC preforms and the sintered preform is not free from pores.

Even liquid phase sintering is not able to provide highly dense SiC preforms and the sintered preform is not free from pores. The average pore size in the solid phase sintered SiC preform has been seen to vary from 1µm to 100µm with the average pore size of 55µm. The porosity and strength varies according to the following equation:Where:= elastic modulus and strength of the porous structure = elastic modulus and strength of the non-porous structure.= Porosity and= constant dependent on pore shapeThe deposition phenomena can be explained as follows. Features of the deposition reaction include: (1) The metal-solution interface which acts as the locus of the deposition process; (2) Kinetics and mechanism of the deposition/infiltration process; (3) Nucleation and growth process of the metal lattice and (4) Structure and properties of the deposits. The reduction of the metal which occurs during the plating process can be generalized.

The reduction of the metal which occurs during the plating process can be generalized. The schematic of the process is shown in figure 2.Figure 2. Schematic of Electrochemical Deposition ProcessTo reduce one mole of given metal, n moles of electrons are required. That is, the total cathodic charge used in the depositionis given by,= gram moles of metal deposited,= Avogadro’s number,= electrical charge perelectronNow(Faraday’s constant), therefore,. Again, we have, wherethe total current and deposited can be written as:is the time of deposition. So total number of moles of metal that can beThe mass of the deposit weight of the metal deposited metal in cc, thickness of the depositcan be obtained by multiplying the above equation with the atomic is the volume of the is the thickness in cm. The.

The. If the density of the deposited metal is is the area of the deposit in cm2 and is given by:,For the specific case of nickel deposition discussed here, the ionic radius of nickel (+2) ions are in the range 0.065-0.073 nm, and its crystal radii vary in the range 0.079-0.087 nm and standard electrode potential for nickel E0 /Ni = -0.25V [7]. The current densities usually range from 540/600 A/m2 and the cathode efficiencies range from 93-97%. By using the equation above, 0.1 mm thick layer of nickel can be deposited on a 2cm cube in 1 hr, which is reasonable for depositing metal ions on pore walls.2+NiExperimentalTwo types of infiltration experiments were carried out viz. Electroless infiltration and Electrolytic infiltration. The experimental procedure followed for both the methods are explained as follows. In preparation for the infiltration experiments, 1 inch cubic samples were made using selective laser sintering.

The experimental procedure followed for both the methods are explained as follows. In preparation for the infiltration experiments, 1 inch cubic samples were made using selective laser sintering. Initial experiments were carried out for the nickel-silicon carbide system. SiC preforms were made using SLS. 180 grit size commercially available SiC powder was mixed with the phenolic binder in the ratio 10:1 and then put inside ball milling jars and allowed to mix for 6-8 hours at moderate speeds. Then the mixed batch of powder was sifted in order to eliminate large sized particles. Prior to that, Differential Scanning Calorimetry (DSC) was carried out on neat phenolic powder to estimate its melting and curing temperatures. The plot obtained is shown in Figure 3. The powder was loaded into the DTM-2000 Sinterstation andisthe processes was carried out at a part bed temperature of 115˚C and laser power of 20W, outline laser power of 4W and slice thickness of about 75µm.

Nitrogen process gas was used to maintain the inert atmosphere inside the machine chamber. After the green parts were obtained, these parts were burnt out in high-temperature furnace at 180˚C/hr to 200˚C, then at 30˚C/hr to 500˚C and then at 300˚C/hr to 800˚C for brown part generation, then to 1400˚C at 250˚C/hr and finally to 1650˚C at 500˚C/hr. After this the furnace was slowly cooled down and the final brown part was obtained.Figure 3. DSC of Phenolic BinderSample Preparation: The following steps were performed in the preparation of the sample for carrying out electrochemical infiltration. Firstly, the cubic SiC preform was cleaned in acetone for 15 minutes and then washed with distilled water. Then the preform is sensitized in 1L aqueous solution containing 10 grams of SnCl2 and 40 ml of HCl, for 15 minutes and then again washed with distilled water.

Then the preform is sensitized in 1L aqueous solution containing 10 grams of SnCl2 and 40 ml of HCl, for 15 minutes and then again washed with distilled water. The preform was activated to open the surface pores by dipping it for 30 minutes in 1L aqueous solution 0.5 grams PdCl2 and 30 ml HCl. After all this the preform was washed and rinsed thoroughly for 10-15 minutes with distilled water [8, 9].Electroless Infiltration: The prepared sample was lowered into a bath containing 45 grams of nickel chloride (source), 8 grams of sodium hypophosphite (reducing agent), 100 grams sodium citrate (complexing agent) and 50 grams ammonium chloride (buffering agent) in 1L of distilled water at pH 9 and temperature of 80ºC. The bath was placed over a hot plate and was stirred using a magnetic stirrer.

The bath was placed over a hot plate and was stirred using a magnetic stirrer. The process was carried out for 90 minutes [8].Electrolytic Infiltration: The process was carried out by attaching a cathode terminal of the electrode to the sample and lowering it in a bath containing 225 grams nickel sulphate, 37.5 grams nickel chloride and 30 grams boric acid in 1L distilled water at pH 3.5. Bath temperature was maintained at 48ºC. The voltage applied between the electrodes was 2.5 V [9].Modified Electrolytic Infiltration: The electrical conductivity of silicon carbide is very low, so a sufficient electronic charge concentration is not achieved uniformly throughout the sample. This inhibits the oxidation of metal ions to form deposit of metals. To overcome this, preforms used as cathode were modified by inserting a conductive graphite electrode at the center of the silicon carbide preform in two configurations as shown in Figure 4.abFigure 4.

(a) Modified Electrolytic Infiltration: Cylindrical Graphite Cathode (b)Modified Electrolytic Infiltration: Rectangular Plate Graphite CathodeThe advantage of this modified process is postulated to be that conductive graphite will attract most of the metal ions through the pores. After the surface of these graphite electrodes are completely saturated with the nickel metal deposits, and the reaction is still continued, the nickel metal will start depositing outwards on the pores and in this way a high part density can be achieved. After carrying out the infiltration experiments, the parts were kept inside a convection oven at 150ºC to dry them for SEM analysis and to remove trapped fluids and gases.Results and DiscussionThe experiments were carried out for both electroless and electrolytic infiltration. The EDS maps showing the distribution of silicon, carbon and nickel is shown in Figure 5.

The EDS maps showing the distribution of silicon, carbon and nickel is shown in Figure 5. The maps were collected at a fractured cross section near the middle of a 1 inch cube infiltrated SiC preform. Figure 5 shows the EDS maps collected using JEOL-JSM 5610 SEM with Oxford Instruments EDS system.Electroless InfiltrationThe results are encouraging because Figure 5 shows that it is possible to infiltrate porous preforms with metal ions using electrochemical techniques by overcoming the surface deposition barrier. The results provide a motivation to carry out other high diffusion rate processes to completely fill the pores and obtain high density functional parts. The EDS maps show that significant amounts of nickel were seen in the preforms without any external diffusion. If an external diffusive force is applied, then higher concentrations of nickel can be seen to occupy the pores.

If an external diffusive force is applied, then higher concentrations of nickel can be seen to occupy the pores. One issue is the removal of undesirable trapped gases and fluids from the bath which mayElectrolytic InfiltrationModified Electrolytic Infiltration (a)Modified Electrolytic Infiltration (b)SiliconCarbonNickelFigure 5. EDS maps showing relative distribution of silicon, carbon and nickel after infiltrationFigure 6. SEM images showing the fractured cross section (a) before (b) after the infiltrationoccupy a large portion of the porous network. Figure 6 shows the SEM images of the preform at a fractured cross-section, for the modified electrolytic infiltration – configuration b.

Figure 6 shows the SEM images of the preform at a fractured cross-section, for the modified electrolytic infiltration – configuration b. It can be seen from these SEM images that modified electrolytic infiltration has been quite effective in depositing the metals in the pores of the SLSed silicon carbide preform.Table 1 shows the weight percentage of nickel deposited in the pores inside the preform in each of the infiltration methods described above. It shows that modifying the process of electrochemical plating by inserting a conducting electrode at the center of the SiC preform leads to an increase in the percentage of nickel deposited in the pores. Higher concentrations of nickel should be achieved by reducing the current density and applying external diffusion. The target composition of nickel to achieve 100% dense part is around 65-70% by weight.Table 1.

The target composition of nickel to achieve 100% dense part is around 65-70% by weight.Table 1. Weight % of Nickel obtained after carrying out each infiltration processProcessAverage % Nickel (by weight)Electroless Infiltration6.4Electrolytic Infiltration3.8Modified Electrolytic Infiltration 19.8Modified Electrolytic Infiltration 218.6It can be seen from Table 1 that the weight percentage of nickel obtained in the preform after infiltration is better in case of configuration b as compared to configuration a. This is because configuration b provided a larger surface area for the ions to be deposited and especially on one side of the electrode which was parallel and facing the anode. The rate of deposition was faster on this side compared to the other side. As a result, more amount of nickel was deposited on the walls of the pores.

The rate of deposition was faster on this side compared to the other side. As a result, more amount of nickel was deposited on the walls of the pores. Hence, orientation of the two electrodes does matter in case ofelectrochemical deposition which can be easily resolved by using two anodes on either side of the cathodic preform.One of the major challenges faced in the current situation relates to the depletion of desired ions into the small pores (where the solution is not assumed to be well-stirred). Initial experiments were carried out just to demonstrate and observe the feasibility of the concept. While the results show incomplete closure of the void network, they are nonetheless encouraging since a significant amount of nickel was deposited deep inside the pores network.

While the results show incomplete closure of the void network, they are nonetheless encouraging since a significant amount of nickel was deposited deep inside the pores network. It is anticipated that external diffusion techniques may well force the electrolyte in and out of the porous preform to improved the extent of deposition.Conclusions and Future WorkInfiltration of SLS silicon carbide preforms with nickel led to encouraging results. The initial stages of work show that the electrochemical infiltration of SLS preforms with metals is highly feasible and has the potential to produce fully dense parts without any undesirable changes. Future work includes exploring other deposition techniques and comparison of those results to these. Ionic diffusion affects the infiltration process to a great extent and hence methods having enhanced diffusion rates such as pressurized diffusion of electrolyte into the preform need to be assessed.

Ionic diffusion affects the infiltration process to a great extent and hence methods having enhanced diffusion rates such as pressurized diffusion of electrolyte into the preform need to be assessed. Other techniques include electrophoretic infiltration in which the charged ions move relative to the electrolyte, through the porous network under the action of applied field. Electroosmotic methods can also be tried where the electrolyte moves with respect to the stationary charged surface. Another important issue is identification of a method for driving the trapped gases and electrolyte out of the porous structure. Infiltration of freeform fabricated parts is a densification method shown to maintain dimensional stability of the part to a high degree and to enable freeform creation of functional parts.References1. H. B. Zhao et al, “Preparation of Palladium composite membranes by modified electroless plating procedure”, Journal of Membrane Science, vol. 142, pp. 147-157. 19982. M.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceElectroforming Process to Additively Manufactured Microscale StructuresKrish Patel, Amelia McNamee, Bishal SilwalDepartment of Mechanical Engineering, Georgia Southern University, Statesboro, GA 30458AbstractElectroforming is a metal forming process that forms parts through electrodeposition. Theoverall study and production of the copper mandrel were conducted by examining the growth and depth of the depositions at macroscopic levels. The study of the plating constant for a copper plate in copper sulfate was performed via the production of copper mandrels. Each mandrel was produced by performing multiple experiments and qualitatively and quantitatively examining the resultant depositions and the initial and final states of the experiment.

Each mandrel was produced by performing multiple experiments and qualitatively and quantitatively examining the resultant depositions and the initial and final states of the experiment. The results were measured based on variations of current and voltage duration in the bath cycle, time duration, solution concentration, and change in mass of both the anode and cathode. The variables such as plating constant and direct current distribution are determined. It seems the rate at which the structure can be fabricated depends on the type of electrolytes used and the parts that are deposited can be scalable.IntroductionElectroforming is a fusion process in which a low voltage charge is passed through a medium in a concentrated premixed electrolytic bath. This breaks down the material so that metallic entities are separated from their original medium. The purpose is to fully separate metallic ions from one substance and apply them to another, as shown in Fig. 1.

The purpose is to fully separate metallic ions from one substance and apply them to another, as shown in Fig. 1. The deposition material is used for the anode, which is positively charged. Once charged, positive ions of the material are transferred to and deposited onto the negatively charged cathode.DC Power supplyCu coatingCathode; - - Cu anodeCuSO 4 elcctrolyt~------· ___ L __ • __ ,_ __ _Figure 1. Electroplating Setup (Kopeliovich, 2013)1697The deposition rate is proportional to the current and time according to Faraday’s Law,which states that the volume of the deposited substance is proportional to the current for a length of time via the plating constant. This can be shown mathematically by the formula:Eq.

This can be shown mathematically by the formula:Eq. 1V=CItin which V is the volume of metal that is plated ( m /Am 3 electrochemical equivalent and density ( for which current was applied (s).mm 3 − s), C is the plating constant based on), I is current (A), and t is the amount of timeThe selection of electrodes and electrolytes depends upon the cell electrode potential. Forcopper-to-copper electrolytic cells (assumes standard 1 molar concentration),Cu(s) | Cu2+(aq) || Cu2+(aq), Cu(s)Ecell = Eox + Ered = -0.34 V + 0.34 V = 0.0 V.Thus it needs to be driven by a power source capable of producing a voltage. Similarly, the reaction could be spontaneous depending upon the reaction as shown below. For Zn in Cu, the reaction is spontaneous and it may not need the additional power source.

Similarly, the reaction could be spontaneous depending upon the reaction as shown below. For Zn in Cu, the reaction is spontaneous and it may not need the additional power source. Similarly, to test the spontaneous voltaic cells, the cell generates 1.10 V.Zn(s) | Zn2+(aq) || Cu2+(aq), Cu(s) Ecell = Eox + Ered = 0.76 V + 0.34 V = 1.10 VThe electrodeposition depends upon the rate at which the positive ions from theelectrolyte enter the crystal lattice and positive ions from the crystal lattice to the electrolyte. The balance between the two aid in the smooth deposition, however one rate may be faster than the other.The objective of this research is to test the feasibility of depositing complex shapes andsizes using the electroforming process. Copper is the most electroplated metal and it has been plated on plastics, die casting parts, jewelery and automotive parts, etc. A FormLabs 2 printer, as shown in Fig.

Copper is the most electroplated metal and it has been plated on plastics, die casting parts, jewelery and automotive parts, etc. A FormLabs 2 printer, as shown in Fig. 2, was used to print precise and accurate models to the desired scale of a honeycomb structure. The printer has the capability of using a stereolithographic resin, which forms polymers when its monomers are photochemically bonded via a laser.1698Figure 2. FormLabs 2 Stereolithography 3D Printer and DesignMethodsOne preliminary study was performed with a home-made electrolyte for 1 hour time with a 1V power supply on a non-castable resin. The results show a deposition height of 134.94 μm as shown in Figure 3.Pilot studies were conducted in order to determine the viability of additive manufacturingvia electroforming. Two different types of electrolytes were used for the experiments. The first experimental setup utilized a home-made solution mixed from copper (II) sulfate crystals in deionized water.

Two different types of electrolytes were used for the experiments. The first experimental setup utilized a home-made solution mixed from copper (II) sulfate crystals in deionized water. The other experimental setup involved using the commercially available Rio Grande Midas solution (Fig. 4). The two electrodes were kept at a fixed distance and the voltage was varied. The initial and final current was measured and the rate of deposition was observed periodically to fill the surface of the honeycomb structure.The honeycomb model shown in Fig. 2 was a prime example of a product that has beenused for simple depositions as the electrodeposition occurs primarily on the base of each hexagonal area due to gravitational effects. In addition, the surface area to volume ratio is high for a honeycomb structure. The dimensions are described in Fig.

In addition, the surface area to volume ratio is high for a honeycomb structure. The dimensions are described in Fig. 2 and the structure is 1 mm in height.A 200 mL beaker was used in the experiment with the copper electroforming solution.The equipment was first cleansed with deionized water and isopropyl alcohol to reduce any residual presence of resin and dirt. The corrosive acid-based solution contained 20% copper sulfate and less than 10% sulfuric acid produced industrially and primarily treats copper metal of1699any sort once diluted. The honeycomb structures were printed with a 1mm thickness and the cathode and anode plates were about 0.7 mm. The plates were placed on opposite sides of the beaker and the required time and current were set. The initial masses of both the anode and cathode, the initial voltage and current, and the volume of solution used were recorded. During runtime, the experiment was monitored for anomalies and records of such were kept.

During runtime, the experiment was monitored for anomalies and records of such were kept. Once the experiment was finished, the final voltage and current were recorded and the anode and cathode were removed from the electrolytic bath and allowed to dry.Figure 3. The surface measurement of the deposited structure using optical profilometer.Once the experiment was set up, electrodeposition commenced. During the electrodeposition process, a current was run through two metallic plates placed in a solution containing metallic ions. These ions were deposited on one of the plates, called the mandrel. Electroforming will take this one step further. Once the plating has finished, the deposited metal will be separated from the mandrel to become its own autonomous part. This part can then be ergonomically shaped for more finely textured geometries and tighter tolerances (Precision1700Eforming LLC).

This part can then be ergonomically shaped for more finely textured geometries and tighter tolerances (Precision1700Eforming LLC). Sintering, the procedure to be used for deposition shaping in future experiments, involves coalescing the metallic ions by applying heat just below the liquefaction point.The construction of the cathode involved trimming the depositing honeycomb structure and using an adhesive to gently apply it to a copper plate. The surrounding external areas were rubber coated with several layers to prevent overgrowth outside the structure. The drying period took approximately 24 hours before experimentation was to begin in order to prevent the breakdown of wet rubber coating. A similar copper plate (anode) was cut out with the same dimensions as the newly created structure but in a more thicker setting to give as much deposition as possible.'l.tCJ,ro f-OWMING COPPER Solutior,Figure 4.

Midas Rio Grande Electroforming SolutionResults and DiscussionThe pilot study results using the home-made electrolytes can be seen in Fig. 5 and Table1. As seen in the results, the deposition time to deposit the high structure was 24 hours. Alterations of the voltage yielded significant changes to the volume and texture of the deposition. The deposition’s thickness is variant based on the voltage and current application but also depends on the placement of both anode and cathode, the volume and concentration of the electrolytic bath, environment, and temperature, etc. It can be observed that the middle two hexagonal structure (1.2 and 1.6 V) has a finer layered structure compare to outer hexagonal structures. The mandrel with 1.2 V does not display as much deposition as expected, while the17012.0 V structure carries more overgrowth on top of the structure, resembling the formation of dendrites.The results of the experiments were promising.

Most of the experiments resulted incopper deposition completely or mostly covering the exposed copper on the cathode. Increased voltage appeared to cause increased deposition rates but a rough surface texture, and growth of dendritic structure, while lower voltage yielded less deposition with a smoother surface. An overall increase in the current could have lead to insufficient ion growth and so it is unable to keep pace with the rate of electron flow which would then cause it to combine with other objects submerged within the solution. The current density is the exponential function of the overpotential [4], thus increasing higher voltage lead to unstable growth in the substrate. The exterior wall of the printed areas was the only visible material and so caused overgrowth within those areas. The surface appears non-uniform.Co noe nt ratio n Voltage (Volt )Initial Current (Amp)Fina l Cu rrent (Amp)Dista ceTi me50g/Ll.

Graph of the volume of copper deposition (cm^3) versus current (A)The overall structures from Fig. 6 shows mild deposition with the commercial electrolyte.Within 1 hr interval. At higher voltage (3.5 V), the mandrel begins to gain mass in the final product as the anode begins to lose mass and begin to rest on the bottom of the beaker. With voltage changes and constant volume, the distance between the plates, and time duration, each trial was run depending on the mass of each plating material and resulting in a corresponding plating constant. The plating constants determined from the data using commercial electrolyte is shown in Fig. 6. As seen in Fig. 6 the relationship is non-linear, this may be due to the faster rate of electron transfer from the anode to the solution than to the cathode, which leads to the disintegration of the anode.The formations within the honeycomb yielded promising results.

Based on the datacollected, the variability in the value of the plating constant indicated a need for more experiments to be performed while holding other factors constant, particularly lower potential.Conclusion and Future work1704Two different types of electroplating solutions were used to deposit copper inside a sacrificial mold. These solutions have a direct impact on the deposition rate. Using a home-made copper sulfate electrolyte, coarse dendritic structures have been grown. A commercial electrolyte solution has been also tested, and finer structure was formed, however, the copper electrode crumbled (disintegrated) into pieces.

ectron Beam Melted (EBM) Co-Cr-Mo Alloy for Orthopaedic Implant Applications R.S. Kircher, A.M. Christensen, K.W. Wurth Medical Modeling, Inc., Golden, CO 80401 Abstract The Electron Beam Melting (EBM) manufacturing process is emerging as an additional method for producing orthopaedic devices in several materials, including Co-Cr-Mo Alloy. This work presents the chemical, microstructural and mechanical properties of several test specimens produced by the EBM process before and after a post‐EBM Heat treatment. Comparisons are made to the properties of Co-Cr-Mo materials used within the orthopaedic implant industry processed by conventional methods such as investment casting and machining from wrought.

Comparisons are made to the properties of Co-Cr-Mo materials used within the orthopaedic implant industry processed by conventional methods such as investment casting and machining from wrought. The results of the work are promising, and demonstrate that EBM produced Co-Cr-Mo material has comparable, and in several cases superior microstructural and mechanical properties to those found in the traditionally-processed materials used today.IntroductionThe last several years have seen many advances towards realizing the goal of additive manufacturing for medical and aerospace applications. The additive manufacturing industry has focused its efforts on producing ‘higher end’ metals including medical grade stainless steels, titanium, titanium alloys and cobalt-chrome alloys. The availability of these materials within the scope of additive manufacturing technology opens the door for many applications in the field.

The availability of these materials within the scope of additive manufacturing technology opens the door for many applications in the field. One of the industry-leading additive metal manufacturing processes is the Electron Beam Melting (EBM) Process. The EBM process is an additive fabrication process, which utilizes a scanning electron beam to melt pre-alloyed metal powder in a layered fashion and build a three- dimensional construct, based on input generated from a computer model.The EBM process is quite different than conventional production techniques, thus it is anticpated that the properties and mirctostructure of EBM produced materials will differ from what is currently seen today. The EBM process intruduces several processing variables which do not exist in currrent manufacturing methods.

The EBM process intruduces several processing variables which do not exist in currrent manufacturing methods. It is important to identify these variables, and come to understand their effect on the metallurgical characteristics of the material.The goal of this paper is to present the metallurgical properties of a particular alloy produced by the EBM process; Co-Cr-Mo. Results of chemical, microstructural, and mechanical evaluations will be presented. The results of these evaluations will be used to compare EBM produced Co-Cr-Mo material to several similar materials used today in the medical industry. In addition, key processing variables that may influence the properties of the material are identified, and their effects on the material are investigated.The Electron Beam Melting ProcessLike many rapid manufacturing techniques, the EBM process creates a physical component from digital CAD models by building the component in a series of layers.

The EBM process starts by distributing a 70 µm layer of fine metal powder onto a steel platform. An electron beam is produced by passing current through a Tungsten filament. The electron beam scans areas of the metal powder layer, in an x-y coordinate system (as defined by the computer model) fully melting the powder in the areas scanned. Once the beam has scanned the appropriate areas, the steel platform is lowered by 70 µm and a new layer of powder is distributed on top of the previously melted layer. This process continues, layer by layer, until a complete part is produced. During processing the entire build chamber maintains a temperatureof approximately 800oC. Once the part has been completed, the build chamber is flooded with He gas to expedite cooling.

During processing the entire build chamber maintains a temperatureof approximately 800oC. Once the part has been completed, the build chamber is flooded with He gas to expedite cooling. A schematic of the EBM process is shown in Figure 1.Figure 1: Schematic of an EBM production system.The use of an electron beam to supply the energy necessary to melt the metal powder mandates that the process be done in a vacuum chamber, which minimizes chemical reactions between the melting metal powder and the surrounding atmosphere. Currently the EBM process is capable of producing parts up to 200x200x200 mm, with a dimensional accuracy of 0.4 mm.Post EBM Heat TreatmentIn addition to evaluating as-produced Co-Cr-Mo samples directly out of the EBM process, samples were also evaluated following a two-stage heat treatment. The heat treatment consisted of a hot isostatic pressing (HIP) cycle and a homogenization treatment.

The heat treatment consisted of a hot isostatic pressing (HIP) cycle and a homogenization treatment. The goals of the heat treatments were to dissolve carbides present in the as-produced material and improve isotropy of the microstructure. Table A outlines the parameters used for each heat treatment cycle.Table A: Heat treatment parametersHeat Treatment CycleTimeTemperaturePressureAtmosphereQuenchHIP4 hrs118525 ksiArgonN/AHomogenization4 hrs1190N/AArgon75oC/MinuteProcedure for Metallurgical EvaluationThe goal of this research was to provide basic matallurgical data from evaluation of EBM produced Co-Cr-Mo alloys, and to compare the results of these evaluations to data in literature and conventionally used standards and specifications.

To accomplish this goal, several test specimens were produced to evaluate the chemical composition, microstructure, and mechanical properties of EBM produced material.Sample Production & Orientation:Mechanical testing was conducted on samples whose gauge-section was both parallel (Z orientation) and perpendicular (XY orientation) to the beam direction. Z orientation samples were produced as cylinders, where as XY samples were produced as large slabs which were sectioned into several specimens.Chemical Composition:Chemical composition analysis was conducted on virgin Co-Cr-Mo alloy powder prior to EBM production. Solid specimens produced from the same powder were analyzed following EBM production to understand changes in chemical composition during solidification. The analyses were conducted in accordance with ASTM E 354 [1].Microstructure:the EBM process were prepared for metallographic examination.

The analyses were conducted in accordance with ASTM E 354 [1].Microstructure:the EBM process were prepared for metallographic examination. The gross microstructure of material was evaluated in two orientations, parallel to the build direction (Z orientation), and horizontal to the build direction (XY orientation). Samples were taken from the mid-section of each specimen, approximately 25mm from the bottom of the build platform. In addition to the EBM produced Co-Cr-Mo alloy, a custom Co-Cr-Mo implant component investment cast in accordance with ASTM F 75[2] was also sectioned and prepared for metallographic examination to allow for a comparison of microstructures.

All samples were evaluated by light microscopy and were electrolytically etched in a 5% HCl solution.Cross-sections of specimens created byMechanical Properties:The mechanical properties of EBM produced Co-Cr-Mo material were evaluated by static tensile testing, rotating bending fatigue testing, Rockwell hardness testing, and facture analysis. The mechanical testing specimens were machined to the appropriate dimensions for each test from the near-net shapes.The static tensile properties of the EBM produced material were evaluated in accordance with ASTM E 8, using the round powder metallurgy specimen geometry (Figure 20 of ASTM E8-08) [3]. Rockwell hardness was measured on cross sections of the grip area on tensile samples following testing. Rotating bending fatigue (RBF) samples were surface polished in the gauge area prior to testing to prevent premature failures. The tests were conducted at a frequency of 100 Hz at room temperature.

Rotating bending fatigue (RBF) samples were surface polished in the gauge area prior to testing to prevent premature failures. The tests were conducted at a frequency of 100 Hz at room temperature. A run-out value of 10 million cycles was chosen.Results and DiscussionTable B contains the chemical composition of the virgin powder prior to production as well as the chemical composition of solid test specimens following EBM production. The results indicate that no significant changes in material composition occur during the melting and re-solidification of the pre-alloyed powder during the EBM process. Of particular note, are the consistent levels of interstitial elements such as O, N, C and H. The high vacuum level in the build chamber of the EBM equipment limits the amount of these gasses available for diffusion into the material during production, allowing for tight control on interstitial elements.

Alloys containing elements with greatly differing melting points have shown a propensity to ‘burn off’ a fraction of the low melting point element during the EBM process. This phenomenon has been observed with a loss of Al during the EBM production of Ti6Al4V alloys [4].This does notappear to be the case for the Co-Cr-Mo alloy system, as all alloying elements remain relatively constant throughout the process.StateTable B: Chemical compositions of Co-Cr-Mo powder and EBM solidified material W wt%Al wt%B wt%C wt%Cr wt%Fe wt%Mn wt%Mo wt%N wt%Ni wt%P wt%Si wt%Ti wt%Powder<0.010<0.0010.2428.290.600.595.740.160.140.0090.370.026<0.01EBM Solidified<0.010<0.0010.2227.690.230.405.960.140.090.0090.380.011<0.01The microstructure of the Co-Cr-Mo alloys resulting from EBM solidification is unique to the process. Micrographs of as-produced EBM Co-Cr-Mo alloy for both XY and Z orientations are shown in Figure 2.

Micrographs of as-produced EBM Co-Cr-Mo alloy for both XY and Z orientations are shown in Figure 2. These micrographs are representative of the microstructure typical seen in EBM produced Co-Cr-Mo alloys. The microstructure consists of columnar grains orientated along the z-axis. A large number of carbides are also present in both micrographs. It is suggested that the combination of high carbon content along with the elevated 800oC chamber temperature during EBM production encourage the precipitation of carbides out of the matrix material.Figure 3 contains micrographs of EBM Co-Cr-Mo alloy for both XY and Z orientations following hot isostatic pressing and a homogenizing heat treatment. The heat treatments result in several significant changes in the material microstructure.

The heat treatments result in several significant changes in the material microstructure. Heat treatment appears to re-dissolve a portion of the carbides back into solution, increase the overall grain size of the material, and most importantly, results in a much more isotropic microstructure.a)b)Figure 2: Micrographs of EBM produced Co-Cr-Mo alloy in the as-produced condition orientated in a) the Z direction and b) the XY direction.Co wt%Bal.Bal.a)b)Figure 3: Micrographs of EBM produced Co-Cr-Mo alloy in the HIP and homogenized condition orientated in a) the Z direction and b) the XY direction.Table C contains the static tensile property results of EBM produced tensile bars in both the as-produced condition and the HIP and heat treated conditions. Test bars from each condition were tested in both the Z orientation as well as the XY orientation.

Test bars from each condition were tested in both the Z orientation as well as the XY orientation. The results indicate that as-produced EBM Co-Cr-Mo material is extremely brittle, especially when stressed along the Z orientation. The low ductility measured in the as-produced material is likely due to large quantity of carbides present in the microstructure. The mechanical property dependence on orientation can be explained by the columnar grains orientated parallel to the beam direction. Heat treatment following EBM production has quite a drastic effect on the mechanical behavior of Co-Cr-Mo material. The 2% offset yield strength is significantly lower and the material exhibits much higher elongation following heat treatment when compared to the as-produced material. The heat-treated samples all exhibited similar mechanical behavior independent of orientation.

The heat-treated samples all exhibited similar mechanical behavior independent of orientation. The average hardness of the heat-treated samples was measured at 33 HRC.Table C: Static Mechanical Properties of EBM produced Co-Cr-Mo material Avg. YS ksi (MPa)Orientation (XY or Z)Number of SamplesAvg. UTS ksi (MPa)Reduction of Area (%)Avg. Elong. (%)StateAs-ProducedXY5104 (717)161 (1110)7.25.0As-ProducedZ5114 (786)126 (869)N/A0.8HIP & HTXY585 (586)166 (1145)24.330HIP & HTZ585 (586)167 (1151)24.030Fracture analysis was conducted on failed samples following tensile testing utilizing a SEM. Figure 3a shows the fracture surface of an as-produced XY orientated tensile sample. The fracture follows the grain boundaries of the columnar grains, and is indicative of a brittle fracture mechanism. Figure 4b shows the fracture surface of a XY orientated tensile that has been heat treated.

Figure 4b shows the fracture surface of a XY orientated tensile that has been heat treated. The fracture can be characterized as intergranular, which typically indicates a brittle fracture with little deformation during failure. This sample, however, exhibited approximately 30% elongation which is rather unique. The microstructures in figure 3 do show a high concentration of carbides at the grain boundaries, which are likely responsible for the intergranular fracture mechanism.a)b)Figure 4: Fracture surfaces of failed static tensile XY orientated samples of EBM produced Co-Cr-Mo alloy in the a) as-produced and b) HIP and homogenized conditions.The fatigue performance of EBM produced Co-Cr-Mo material is summarized as an S-N curve in Figure 5. Two groups of specimens were tested; samples orientated in both XY direction and the Z direction. Both sample groups underwent HIP and homogenization treatments prior to testing.

Two groups of specimens were tested; samples orientated in both XY direction and the Z direction. Both sample groups underwent HIP and homogenization treatments prior to testing. The number of samples in each group is not sufficient to accurately predict the 10 million cycle stress level, however it appears the value will be between 400 and 500 MPa. The data also indicates that samples tested in the Z direction have a slightly lower fatigue limit.

The data also indicates that samples tested in the Z direction have a slightly lower fatigue limit. This may be a consequence of the layering in the Z direction which could provide crack initiation sites.750700CoCrMoXY Orientation) a P M650600Z Orientation( x a m σσσσ55050045040035010,000100,0001,000,00010,000,000Cycles (Nf)Figure 5: S-N curve for EBM produced C-Cr-Mo material tested in rotating bending fatigue.If EBM produced Co-Cr-Mo material is going to be used in the medical field it is important to understand how all of the characteristics presented in this study compare to the characteristics of the Co-Cr-Mo materials currently used in the medical industry. Table Dsummarizes the most commonly used Co-Cr-Mo materials in the medical industry today [2]. The main difference between these alloys is the method of production; forging, casting or wrought alloys.

The main difference between these alloys is the method of production; forging, casting or wrought alloys. Each alloy has slightly different chemical composition requirements set at the necessary levels to enable proper manufacture and result in the desired mechanical performance. The powder used in this study was produced with the intention of meeting the ASTM F-75 chemical composition requirements. Table E compares the chemical composition of the solid EBM Co-Cr-Mo alloy with that of each of the standard materials listed in Table D. The ASTM F799 and F1537 standards designate three (3) separate alloys; 1) Low carbon, 2) High Carbon and 3) Dispersion Strengthened.

The EBM produced material meets the chemical requirements of ASTM F 75 as well as high carbon (Alloy 2) material.Table D: Conventional standards for Co-Cr-Mo alloys in the Medical IndustrySpecification DesignationTitleASTM F 75Standard Specification for Cobalt-28 Chromium-6 Molybdenum Alloy Castings and Casting Alloys for Surgical ImplantsASTM F 799Standard Specification for Cobalt-28 Chromium-6 Molybdenum Alloy Forgings for Surgical ImplantsASTM F 1537Standard Specification for Wrought Cobalt-28 Chromium-6 Molybdenum Alloys for Surgical ImplantsTable E: Chemical composition comparison of EBM produced material and conventional Co-Cr-Mo AlloysMaterialAlloyAl wt%B wt%C wt%Cr wt%Fe wt%Mn wt%Mo wt%N wt%Ni wt%P wt%Si wt%Ti wt%W wt%EBM Solidified……<0.010<0.0010.2227.690.230.405.960.140.090.0090.380.011<0.01ASTM F 75ASTM F 799 & F 1537 ASTM F 799 & F 1537 ASTM F 799 & F 1537…….123<0.10…………0.30- 1.00<0.010………………<0.35<0.140.15- 0.35<0.1427.0- 30.0 26.0- 30.0 26.0- 30.0 26.0- 30.0<0.75<0.75<0.75<0.75<1.0<1.0<1.0<1.05.0- 7.0 5.0- 7.0 5.0- 7.0 5.0- 7.0<0.25<0.25<0.25<0.25<0.50<1.0<1.0<1.0<0.020………………<1.0<1.0<1.0<1.0<0.10………………<0.20………………Figure 6 compares the mechanical properties of EBM produced Co-Cr-Mo material following HIP and heat treatment with the mechanical requirements of the alloys listed in table D.

There are several important conclusions that can be drawn from the comparison. The elongation and reduction of area properties of the EBM produced materials exceed the requirements for all conventionally produced materials. The EBM produced material also meets or exceeds all mechanical strength requirements for both ASTM F75 and ASTM F1537 Annealed material.Although none of the ASTM standards for implantable Co-Cr-Mo alloys require a specific microstructure, it useful to compare the the microstructure of EBM produced materials to that of a conventionally produced implantable device. Figure 7 is a optical light micrograph of an investment cast commercially available implantable Co-Cr-Mo device which was obtained and sectioned for microstructural evaluation. It is obvious that the microstructure of this implant is quite different than the microstructures presented in figure 3.

It is obvious that the microstructure of this implant is quite different than the microstructures presented in figure 3. The microstructure in figure 7 is what one would expect from an investment cast material that experiences a slow solification rate [5].

EBM produced material solidifies very quickly after melting, and experiences anLa wt%N/A……......……0.03- 0.20Co wt%Bal.Bal.Bal.Bal.Bal.entirely different thermal history through out its production and subesequent heat treatments, resulting in an entirely different microstructure.180160EBM Solidified & Heat treated140ASTM F 75120ASTM F 1537 Annealed100ASTM F 1537 Hot Worked80ASTM F 1537 Warm Worked60ASTM F 79940200Yield StrengthUTSReduction of Area (%)Elongation (%)Hardness (HRC)Figure 6: Comparison of mechanical properties of conventional medical grade Co-Cr-Mo alloys and EBM produced Co-Cr-Mo alloy.Figure 7: Microstructure of a commercially available investment cast Co-Cr-Mo implant.Summary and ConclusionsThe purpose of this research was to investigate whether EBM produced Co-Cr-Mo alloys exhibit adequate chemical and mechanical properties to be considered acceptable alloys for surgical implant applications.

The results demonstrate that as-produced Co-Cr-Mo alloy is quite brittle and exhibits anisotropic mechanical behavior. For these reasons, as-produced Co-Cr-Mo material needs to undergo a two-stage heat treatment in order to satisfy the requirements of standard Co-Cr-Mo medical grade standards. The heat treatment has the effect of dissolving carbides which precipitate out of the matrix during EBM solidification, and removing the anisotropic behavior from the material. Although the microstructure of EBM produced Co-Cr-Mo alloy is unique when compared to material manufactured by conventional means, its mechanical performance exceeds or meets the requirements for both ASTM F75 material and annealed ASTM F1537 grade material.1. ASTM Standard: E 354 Standard Test Methods for Chemical Analysis of High- Temperature, Electrical, Magnetic, and Other Similar Iron, Nickel, and Cobalt Alloys, 2006.1.

ECTRON BEAM SOLID FREEFORM FABRICATION OF METAL PARTS v. R. Dave*, J. E. Matz:j:, T. W. Eagar:j:ABSTRACTA novel method for near-net-shape fabrication of large, complex metal parts without the needfor specific tooling is described. Parts are built up in a layer-by-Iayer fashion by feeding raw material in wire form into a melt pool which is maintained with an electron beam. Several different alloys have been successfully deposited using a low-energy electron beam. A model has been developed which permits determination of the melt depth and vaporization-limited power input as functions of beam energy and travel velocity. Processing parameters obtained from experiments with low-energy electron beams can be used to predict melting behavior under the influence of higher energy beams.

Processing parameters obtained from experiments with low-energy electron beams can be used to predict melting behavior under the influence of higher energy beams. The process is being investigated as a possible means for manufacturing compositionally-graded components.INTRODUCTIONThere are very substantial military and economic incentives to return naval and commercialships to service as soon as possible after a failure. Unfortunately, the casting, forging and machining processescurrently used to make large replacement drive components such as shafting, couplings, bearings, and propellers, can require lead times of several months. This necessitates maintenance of a large inventory of spares for every vessel in service. The manufacturing costs and warehousing fees for these spares can be a very large portion of the total operating expense for a ship.

The manufacturing costs and warehousing fees for these spares can be a very large portion of the total operating expense for a ship. Manufacture of parts for initial construction can also be costly for short production runs (i.e., if only a single ship of a given design is being built).One of the most time-consuming and expensive steps in producing components via aconventional route is machining of the part-specific tooling required for casting and forging. Mold patterns, permanent molds, and collapsible cores are required for each casting. Forging demands even more tooling; in some cases, a different die for each reduction step is needed. Moreover, since the dies must withstand high stresses and temperatures, they are generally made from alloys which are very hard at room temperature. This can result in extremely low machining productivity.

This can result in extremely low machining productivity. The inherent limitations of the casting and forging processes in producing certain geometric contours can also result in considerable machining of the fmal part.Conventional processing is also somewhat less than optimal from the standpoint of the mechanical properties obtained. The main reason for this is inhomogeneity caused by solute rejection and subsequent convection during solidification. The ensuing macro-segregation will occur over a distance comparable to the dimensions of the casting. It can theoretically be eliminated by a homogenization anneal, but the time required would be excessive due to the long diffusion distance. Forging has little effect on the segregation, and it can also introduce inhomogeneity through non-uniform deformation.A number of processes have been developed over the past decade which attempt to reduce reliance on part-specific tooling and machining.

The most prominent of these Solid Freeform Fabrication (SFF) techniques are: 3-D printing (3DP), Selective Laser Sintering (SLS), and spray forming. All of these processes have certain advantages and disadvantages, and each is superior to the others for certain types of fabrication problems. For example, 3DP and SLS are capable of very high dimensional accuracy for net-shape manufacturing, but productivity suffers due to the needUTC/Pratt & Whitney, East Hartford, CT :j:MIT Materials Science and Engineering Department, Cambridge, MA64for powder spreading and leveling between each layer. On the other hand, spray fonning can produce large parts at high deposition rates, but only simple shapes can be created.ELECTRON BEAM SOLID FREEFORM FABRICATIONAnother approach exists which provides a combination of capabilities uniquely suited to thetask of quickly fabricating large components to near-net-shape from metals of engineering significance.

In this process, called Electron Beam Solid Freefonn Fabrication (EBSFF), parts are built up in a layer-by-Iayer fashion as in other SFF techniques. However, the similarity ends there; the addition of the metal comprising each layer is accomplished by feeding wire into a melt pool maintained on the surface of the previous layer by an electron beam (shown in Figure 1 for an annular part geometry). A positioning system pennits controlled motion of the melt pool, and hence the location where the deposition is taking place. Successive beads of deposited metal are placed adjacent to each other to fonn a layer. In this respect, the process is somewhat similar to multi-pass welding.The EBSFF process is very flexible. Nearly any material which can be melted at reasonablepower levels and is obtainable in wire fonn can be used as feedstock.

Nearly any material which can be melted at reasonablepower levels and is obtainable in wire fonn can be used as feedstock. The latter requirement effectively limits consideration to metals, but ceramic powders may be incorporated via cored wire if they can be made to be spontaneously wet by the sheath metal. This approach may be of use for manufacturing cennet cutting tools and other composites. In any event, near tenn research will focus on alloys of interest for nautical service: stainless steels, rotor steels, bronzes, brasses, and aluminum and titanium alloys.The versatility of the process is further illustrated by the wide variety of part geometries whichcan be fonned. Parts can be built up on a flat surface by horizontal translation relative to the electron beam, or annular parts can be fashioned by deposition onto a rotating cylindrical substrate. The part can be re-oriented to produce features on virtually any of its surfaces.

The part can be re-oriented to produce features on virtually any of its surfaces. Prudent choice of the substrate and orientation allows internal channels to be created. Of course, some geometrical constructions cannot be fonned, but this is not expected to be a problem for most parts.The limitations of the process are of great interest for comparison to conventional processing and to other SFF methods. The maximum component size is expected to be limited by cracking caused by tensile residual stresses in the part. Temperature gradients are set up in the part as a consequence of the heat input associated with the deposition process. Differential thennal contraction during cooling after processing leads to the residual stresses. Careful control of the post-processing cooling conditions may mitigate this problem, but a better solution is possible due to a unique feature of electron and other high-energy-density beams.

Careful control of the post-processing cooling conditions may mitigate this problem, but a better solution is possible due to a unique feature of electron and other high-energy-density beams. If designed for the purpose, the beams can produce pulses of microsecond duration. The ensuing rapid local thennal expansion leads to shock wave propagation throughout the part. Localized plastic defonnation occurs at the shock wave front as it traverses the part and this greatly reduces the elastic stresses. If this is done periodically during processing and while the part is being cooled from the processing temperature, the residual stresses can be reduced to levels below that which will cause cracking. There is no significant macroscopic dimensional change associated with the treatment.The maximum productivity (deposition rate) attainable in the process is expected to be limited by a combination of vaporization from the surface of the melt pool and heat transfer from the part.

Methods for reducing vaporization while maintaining high heat input will be dealt with in the modeling section below. After the heat energy is used for deposition, it must be removed from the part as quickly as possible. If the bulk temperature of the part becomes too high, it may lose its shape or its fine solidification microstructure may coarsen. Active cooling of one or more of the part's surfaces will be necessary to limit this temperature rise. For making large parts from alloys with low thennal conductivity, even this may not be adequate and an atmospheric high-energy electron beam may have to be used in conjunction with convective cooling by a liquid. The extreme65temperature gradients caused by active cooling will make provision for some relief during processing all the more imperative.of shock stressIt is not expected that the process will be capable of producing true net-shape parts.

The spatialresolution of the technique will be limited to a distance approximately equal to the diameter of the feedstock wire. Since relatively large wires can and should be used to attain the high production rates necessary to economically produce large components, dimensional accuracy will be low. Furthermore, the as-produced surface finish is unacceptable for most purposes since there is usually some waviness arising from the one-dimensional nature of the deposition process. This means that some post-machining will always be necessary to achieve the required dimensional accuracy and surface fmish. However, it is important to note that this machining will involve the removal of a relatively small volume of material and so will not contribute significantly to the total time required for fabrication.

Post-processing of castings (gating system removal, weld repair of defects) and forgings (flash removal) can be far more costly and time-consuming.The mechanical properties of material produced by this method are expected to be superior to those of cast parts of the same material, but inferior to those of forgings. The improvement over castings arises because only a small volume of metal is solidifying at anyone time. This leads to a higher local solidification rate and a finer and more uniform microstructure. Properties are also expected to be more homogeneous since the extent of segregation is limited to less than the size of the bead cross-section. It is feasible to eliminate segregation of this degree via a homogenization treatment.

It is feasible to eliminate segregation of this degree via a homogenization treatment. Plastic deformation comparable to that in forging is not intrinsic to the deposition process, so parts produced by this process will not attain the strength levels of forged parts.Another impediment to commercialization of the EBSFF process is the large capital investment required. For low-energy EBSFF, the electron beam generation equipment is not expensive, but a vacuum chamber large enough to contain the part will be required. For high-energy EBSFF, the electron accelerator itself is costly, and heavy radiation shielding is required to protect personnel and equipment from the hard x-rays produced. On the other hand, a high-energy (>500 keV) beam can propagate in air with minimal attenuation, so vacuum equipment is not necessary.have been made with several materials using a commercial low-energy (120 keV)electron beam welding apparatus. A layer of 308L stainless steel 0.5 in.

A layer of 308L stainless steel 0.5 in. thick and 0.75 in. long was successfully deposited onto a 304L stainless steel mandrel with an outside diameter of 4.5 in. Figure 2 is a micrograph of a radial section of the deposit; the individual beads can easily be distinguished due to their different solidification environments. A commercially pure titanium deposit (0.3 in. thick, 1.0 in. long) was made in the same manner. This specimen exhibited some porosity at the interface between the substrate and first layer of weld beads; this was probably caused by inadequate cleaning of the substrate prior to deposition.The EBSFF process is not only an alternative way of making parts from traditional materials,but can also be used to make materials which are difficult to process conventionally. It is extremely suitable for making parts with composition gradients.

It is extremely suitable for making parts with composition gradients. By changing the feed wire at appropriate times during deposition of a part, one-dimensional composition variation can easily be achieved. Continuous composition variation is accomplished by using powder-cored wire as feedstock; an appropriate metering device can be used during the fabrication of the cored wire to continuously vary the composition of the powder in the core. The process also provides a very controllable way of reactively forming intermetallic parts, although multiple wire feeders will probably be needed to do this economically.PROCESS MQDELINGIt was mentioned above that vaporization from the surface of the melt pool is expected to be amajor factor limiting the productivity achievable by the EBSFF process. Excessive vaporization not66reducesamount of thennal energy availablethe melt pool for deposition, it also lowers'It"t"Ilott:l>rtoi yield.

Excessive vaporization not66reducesamount of thennal energy availablethe melt pool for deposition, it also lowers'It"t"Ilott:l>rtoi yield. Furthennore, if any alloy constituents have very dissimilar vapor pressures, significant differences between the composition of the feedstock material and the part may result. Damage to the equipment for electron beam generation, vacuum production, wire feeding and part also occur. It is therefore of the utmost importance to be able to accurately predict positioning the amountvaporization for various operating conditions.a given alloy, the quantity of vaporization will be largely determined by the temperature at surface of the melt pool. This will be determined by the thennal balance between energy addedby the beam and heat conducted away into the part.

This will be determined by the thennal balance between energy addedby the beam and heat conducted away into the part. Since electrons are known to dissipate their kinetic energy within a volume of material below the surface instead of at the surface, a transfer model which can simulate a volumetric heat source is required. Additionally, the model must be capable of describing the effects of motion of the heat source. Dave [1] has expanded the function model of Eagar and Tsai [2] for moving distributed surface heat sources to alsoencompass volumetric heat sources.An effort was made to take into account all factors relating to the energy deposition and heattransfer, but have a Gaussian distribution of energies. The beam power is of interest for operational reasons, but the power absorbed by the material is the relevant quantity for heat transfer calculations.

The beam power is of interest for operational reasons, but the power absorbed by the material is the relevant quantity for heat transfer calculations. Consequently, the model results are displayed in tenns of the absorbed power; the beam power can obtained through application of backscattering coefficients appropriate for the material involved.some cases this could not be done. Electrons in the incident beam were assumed toLimitations of the Green's function technique prevent radiation from being taken into account directly, but its contribution was estimated by assuming the surface of the pool to temperature. In the interest of simplicity, convection in the melt pool was assumed not to occur, and the physical properties of the liquid and solid were assumed to be the same and constant with temperature.

In the interest of simplicity, convection in the melt pool was assumed not to occur, and the physical properties of the liquid and solid were assumed to be the same and constant with temperature. While the numerical accuracy of the results produced by the model will clearly compromised by these assumptions, qualitative conclusions drawn from the results are still expected toat the meltingvalid.Low-energy electrons are expected to have very limited penetration into the interiora'It"t"IlO,tt:l>1I""O and so they will behave as surface heat sources. Comparison of the models for surface and volumetric Increasing the beam energy causes a proportionally greater amount of energy tosources confrrms that this is true for beam energies below about 200 keV.deposited belowmaterials' surface. This enables a greater depth of material to be melted without increasing the 304 increases power required tosurface temperature (and hence, stainless steel.

This enables a greater depth of material to be melted without increasing the 304 increases power required tosurface temperature (and hence, stainless steel. However, the power input required to obtain a given amount This is illustrated in Figure 4, which shows the amountamount of vaporization), as shown3 forbeaminitiate melting as a function of beam energy. For a given beam energy, absorbed power l'M"t:l>'Cltt:l>1I'" than those required to initiate melting, but lower than those which will cause excessive 'H01r'\n.... r:7or.n."l can be used for deposition.

r:7or.n."l can be used for deposition. This defines a processing space as shownacceptable vapor pressure has been arbitrarily set to one atmosphere practicalwill probably be much lower.":u.,,vu"v'" • ...,"".....I".'t"r('}'.."r variable affecting the attainable productivity istravel velocity.to a stationary beam, a rnt:l>ltll',O" or vaporization, since it is always encountering fresh (cooler) .........,..v"''''...,...beam will have to have higher power to causesameprc.ce~)srnlgcan therefore alsodefmed with respect to the travel velocityplots showing the processing space can be used to obtain the total power available for melting of the feed wire) by subtracting the power required to Jl,luriI""..·.,"'C'.'hr'n"'"vU.llJ';::;'absorbed power. subtracted. This canthe power required for maintenancecalculated for a given absorbed power a melt pool cross-sectional.. ,,"" .r-.p'lr..

subtracted. This canthe power required for maintenancecalculated for a given absorbed power a melt pool cross-sectional.. ,,"" .r-.p'lr.. , withadditional assumption67rectangular cross-section, the power required to maintain the of the total power available for melting and deposition. The deposition rate and the deposition energy efficiency can be calculated from the deposition power. The higher absorbed power levels possible with higher energy beams cause an increase in the fonner quantity (the specific deposition rate is roughly constant at 2.2 Ib/hr/kW), but a decrease in the latter one (dropping from 0.41 for 100 keV electron beams to 0.39-0.40 for 1.0 MeV beams). That is, higher productivity can be obtained by going to higher beam energies, but only at the cost of lower energy efficiency.pool is typically about 10-15%It is useful to compare these quantities to those obtained with deposition accomplished with a variety of other heat sources.

Obviously, lasers, causing only surface heating, will exhibit behavior very similar to that of low energy electron beams. The choice between these two heat sources will largely depend on a comparison of the emissivity of the surface for the laser and the backscattering coefficients of the material for the electron beam. Gas-metal arc welding (GMAW) is another competitive process. Calculations made on the basis of production welding schedules indicate that both deposition rate per unit absorbed power (1.0-1.5 Ib/hr/kW) and deposition energy efficiency (0.20-0.26) are approximately half that for electron beams, even low energy ones. This is probably due to the additional power required to produce and maintain the plasma in the arc.The model's validity was checked by using it to calculate the power required for deposition ofthe 308L stainless steel sample described in the previous section, and comparing it to the actual power used.

The model result was 1141 Wand the actual value was 1248 W, a difference of approximately 10%. This is remarkably good agreement, and provides very resounding evidence for the applicability of the model to EBSFF.SUMMARyThe EBSFF process is a viable way of making large metal components to near-net-shape much more quickly than the established methods (casting, forging, machining). This is confrrmed by trial deposits which have been made from several relevant alloys. It may also be useful for making parts from advanced materials which are difficult to make by conventional means, such as particulate reinforced composites, intennetallic compounds, and materials with composition gradients. However, high capital cost is a significant drawback relative to other SFF processes.A model for the heat transfer from a moving volumetric heat source has been developed using a Green's function solution for the heat conduction equation.

This model is appropriate for modeling the interaction of high energy electron beams with metals. A number of parameters of technological interest relating to deposition with high energy beams can be obtained from the model, such as the vaporization-limited maximum deposition rate and the energy efficiency of the process. Comparison with competitive processes, such as laser and arc-welding-based deposition, indicates that high energy electron beams are expected to be capable of higher production rates and energy efficiency.The authors wish to gratefully acknowledge Diego Penta for his assistance figures for this paper. This work was supported by the Office Naval Research.REFERENCES1.V. Institute of Technology, 1995.Dave, "Electron Beam-assisted Materials Fabrication," Ph.D. Thesis, MassachusettsW. Eagar and N. S. Tsai, "Temperature Fields Produced by TravelingHeatSources," The Welding Journal, 62 (12), 1983, pp. 346s-355s.68--t,';, -_.~......--Figure 1.

ectronics Integration in Conformal Substrates Fabricated with Additive Layered Manufacturing Sylvia Castillo, Dan Muse, Frank Medina, Eric MacDonald and Ryan Wicker The University of Texas at El Paso, W.M. Keck Center for 3D Innovation, El Paso, Texas 79968A three-dimensional (3D) accelerometer sensor system with microprocessor control was fabricated using a previously developed integrated layered manufacturing system that combines conductive ink dispensing with stereolithography (SL). The electronics are integrated into a conformal substrate that is press-fit into a helmet for the purpose of detecting Traumatic Head Injury (THI) when an excessive acceleration to the head is measured. Applications include monitoring the health of soldiers or athletes. Traditional fabrication of electronics is implemented with a 2 dimensional printed circuit board (PCB), which are not well suited for rugged installations in curved locations such as the interior of a helmet.

The advantage of layered manufacturing for the integration of electronics is the ability to fabricate in a conformal substrate - conforming to the curved, complex, and often flexible shapes dictated by the human body.Keywords: rapid prototyping; stereolithography; direct-print; hybrid manufacturing, conformal electronics, accelerometer, three-dimensional electronicsIntroductionThe capability of designing and fabricating electronic systems, which are both conformal and can easily be customized at the unit-level will provide a breakthrough in the fabrication of bio-medical engineering devices. Furthermore, mechanical systems that require integrated electronics in structural components would stand to benefit as well (e.g. sensors and microprocessor systems embedded in the nose cone of an unmanned aerial vehicle). Human anatomy requires medical devices to have non-orthogonal, curved surfaces that may be either flexible or stiff depending on the application (e.g.

Human anatomy requires medical devices to have non-orthogonal, curved surfaces that may be either flexible or stiff depending on the application (e.g. bandages need to be flexible and stretchable while prosthesis require stiffness). As body shapes vary from person to person, the capability of customizing shapes of bio-medical devices is particularly important, as the strategy of one-size- fits-all is not well suited for these applications.Implementing electronics systems that are conformal with curved and complex surfaces is difficult if not impossible with traditional fabrication techniques, which require stiff, two dimensional printed circuit boards (PCB). Flexible copper based fabrication is currently available commercially providing conformance but cannot be simultaneously stiff. Consequently these systems are susceptible to reliability problems if bent or stretch repeatedly.

Consequently these systems are susceptible to reliability problems if bent or stretch repeatedly. The integration of additive layered manufacturing with micro-dispensing can provide shapes of arbitrary and complex form including miniature cavities for electronics components and conductive traces for electrical interconnect between components.The fabrication freedom introduced by Additive Layered Manufacturing (ALM) techniques such as stereolithography (SL), ultrasonic consolidation (UC), and fused deposition modeling (FDM) have only recently been explored in the context of electronics integration. Advanced dispensing processes have been integrated into these systems allowing for the introduction of curable conductive inks and epoxies to serve as electrical interconnect within the ALM dielectricstructures.

This paper describes a process that provides a novel approach for the fabrication of stiff conformal structures with integrated electronics and describes a prototype demonstration – an accelerometer sensor with a microcontroller and wireless reporting capability. The intent of the demonstration is to act as an insert to the helmet of a soldier or athlete to detect the Traumatic Head Injury and report detection as a concussion can result in the injured person being unaware of the problem.Previous WorkThough reports in literature are sparse, a growing number of researchers have shown interest in the capability of fabricating 3D and conformal electronics using Additive Layered Manufacturing (ALM).

The combination of Direct Printing (DP) of conductive inks onto SFF structures was introduced by Palmer et al., (2004) and expanded in Medina et al., (2005) and Lopes et al., (2006) in which simple circuits were implemented to demonstrate functionality by integrating a dispensing system into an SL machine using three-dimensional linear stages with a dispensing head. This approach included a demonstration of a simple prototype temperature sensor with nine components including a 555-timer chip. Periard et al., (2007) demonstrated a similar circuit as well as several clever electro-mechanical applications all created by an open- source fabrication system. Navarrete et al., (2007) describes improvements to using DP on SFF substrates by introducing channels into the substrate for the conductive material in order to provide delineation of the electrical lines and allow for the reduction of line pitch, width and spacing while reducing the possibility of line-to-line shorting.

Line spacing was thus controlled by the precision of the SL fabrication (e.g. laser beam size) rather than the dispensing process. Furthermore, the demonstration of this technique included not only digital electronics (e.g. PIC processor and GPS chip set) but also included high frequency (RF) functionality (e.g. antenna conductors). The electronics were implemented in a shape of a camouflaged rock to highlight the possibility of creating intricately detailed and arbitrary-formed devices made possible by SFF. All of the reported circuits to date required only the use of a single plane of routing (e.g. no crossing conductors) although the concept of multiple planes with vertical interconnects was the obvious next step. Advancements in routing of DP electrical interconnect integrated into ALM structures was described in Palmer et al., (2004) and is considered seminal for the work described in this report.

Advancements in routing of DP electrical interconnect integrated into ALM structures was described in Palmer et al., (2004) and is considered seminal for the work described in this report. General advancements in dispensing techniques that may be well suited for integration into ALM structures was described in Church et al., (2005) in which conductive lines were drawn onto glass substrates in order to create wireless sensor systems. The described proprietary pumping system provided precise lines with widths as small as 25 microns while drawing at speeds as high as 250 mm per second. This technology is capable of more than planar processing and can dispense conductive or dielectric materials onto three-dimensional conformal structures (e.g. drawing an antenna conductor onto a soldier’s helmet).

drawing an antenna conductor onto a soldier’s helmet). The integration of this advanced printing technology with our ALM fabrication is the subject of on-going collaborative work and will provide for promising improvements to routing density and speed of fabrication of next generation ALM-integrated electronics. Moreover, this technology demonstrated the possibility of printing not only the conductive interconnect but also passive electrical components such as capacitors, inductors and resistors, and consequently may provide for further miniaturization capability. Arnold et al., (2007) described a technique referred to as Laser Induced Forward Transfer (LIFT) that allows for the deposition of very thin lines in a variety of materials including copper. A timing circuit similar to the ones described previously was demonstratedwith bare silicon die and unpackaged surface mount passive.

A timing circuit similar to the ones described previously was demonstratedwith bare silicon die and unpackaged surface mount passive. In addition to highly precise conductor deposition, the paper describes the possibility of fabricating batteries. This work did not include ALM substrates and is limited to two-dimensional deposition. Fig. 1 illustrates three generations of a 3D off-axis placement and routing of a magnetometer system, which included a microprocessor, LEDs, a DC connector and three orthogonally placed magnetic hall effect sensors (de Nava et. al. 2008). The first generation was almost entirely built on one surface of the structure as designing and routing on two dimensions is supported with current electronics CAD programs and is significantly simplified as a result. The top face contained a majority of the circuits and routing. Included on the top face were two of the three hall effect sensors (one for the X direction and the other for Y).

The top face contained a majority of the circuits and routing. Included on the top face were two of the three hall effect sensors (one for the X direction and the other for Y). The third sensor was placed alone on an alternate side to provide the Z axis. The second and third generations were successively improved and miniaturized as all six faces of the structure were utilized and the was microprocessor interior.

The second and third generations were successively improved and miniaturized as all six faces of the structure were utilized and the was microprocessor interior. embedded These final versions included significantly improved volume utilization, were significantly more difficult to design as CAD software does not exist the design freedom required for a 3D design in terms of placement and routing All though these demonstrations did illustrate the possibility of 3D fabrication with off-axis placement and routing, all surfaces were flat and all angles were at 90 degrees.inthebutFigure 1 – Three Generations of a Three-Axis Magnetic Flux Sensor Systemthat providesIn the current work, a novel approach of using ALM techniques coupled with conductor dispensing to provide a dielectric substrate that is conformal, stiff and three-dimensional is described.

A similar sensor fabricated with traditional flexible PCB techniques would have compromised reliability and also result in bunching to conform to the interior of the helmet. Furthermore, the system includes wireless capability with a ink-based antenna. The 3D design freedom could provide for enhanced RF capability as the antenna design is not limited to just a 2D or conformal curved surface, however in this report, the Antenna was applied directly to the helmet insert surface.Conformal Electronics Design Methodology and FabricationThe design of the Accelerometer System Helmet Insert began with modeling of the geometry inside the target helmet. A silicone mold was made from inside of the helmet, measurements were taken, and a 3D model of the basic shape, shown in Fig. 2(a), was created in SolidWorks. An extruded circular plate was then created with some reference art included for later features. Using the deform feature in SolidWorks, this plate was shaped to match the helmet form.

An extruded circular plate was then created with some reference art included for later features. Using the deform feature in SolidWorks, this plate was shaped to match the helmet form. Figure 2(b) and (c) show the plate before and after deformation. Holes were cut into this deformed(a)(b)(c)Figure 2 – Deformation Processshape to correspond to the mounting holes on the target helmet. These holes also served as the general boundaries for the design. Using arcs to define the outer perimeter of the insert, excess material was cut from the model. Fillets were used at the mounting holes to ensure mechanical integrity and give a purposeful appearance. The overall contour of the insert shown in Fig. 3(a) was produced in stereo lithography to ensure proper fit and alignment in the target helmet.Once the overall form of the insert was confirmed, the electrical design could begin.

3(a) was produced in stereo lithography to ensure proper fit and alignment in the target helmet.Once the overall form of the insert was confirmed, the electrical design could begin. The two primary integrated circuits used in the design (an accelerometer and a microcontroller with integrated transmitter) use 26mil fine pitch leads and require precise geometry for effective implementation in SLA. As such, a small portion of the insert required a flat surface from which to develop the circuit geometry. Figure 3(b) shows an elliptical area which was cut in the top of the design to provide the flat surface. Also shown are the two cavities where the ICs will be placed. Using SolidWorks tools such as construction lines, offset entities, and extruded cuts, the circuit interconnections were created on the flat surface as shown in Fig. 3(c). The channels and thru-holes provide a consistent path for the dispensing of conductive silver ink. Also shown in Fig.

3(c). The channels and thru-holes provide a consistent path for the dispensing of conductive silver ink. Also shown in Fig. 3(c) are extruded cavities for several discrete components such as capacitors, resistors, a crystal and LED.(a)(b)(c)Figure 3 – Overall Shape and circuit developmentTo reduce duplicate design work, reusable models were created based on the pin and package geometry for both ICs (Fig. 4(a)). These models were imported into the design, creating an assembly. Figure 4(b) and (c) shows these components before and after insertion into the design.(a)(b)(c)Figure 4 – Reusable component modelsTo complete the design, a rectangular loop antenna along with a battery cavity and ground plane (under-side of design) were extruded. The antenna, though not conformal, was extruded from a plane tangential to a point on the curved surface, thus keeping the antenna within the contour of the design.

The antenna, though not conformal, was extruded from a plane tangential to a point on the curved surface, thus keeping the antenna within the contour of the design. A flat antenna surface was used to maintain a symmetrical radiation pattern giving optimum antenna performance. The battery cavity was designed to hold a standard Lithium coin cell housing and is connected to the insert’s electronics by way of a channel on the top surface of the design and an extruded guide on the bottom surface. Ultimately, our goal is to use lithium polymer technology, which will allow our battery to not only be rechargeable, but also be formed to match the contour of the insert. A more ambitious goal is to utilize Direct Print technology to print a conformal battery system onto the insert. The ground plane is the elliptical area cut directly beneath the circuit on the bottom of the design as shown in Fig. 5.

The ground plane is the elliptical area cut directly beneath the circuit on the bottom of the design as shown in Fig. 5. Its purpose is two-fold: to provide convenient ground connection to several points of the circuit (through holes to the surface); and to improve the radio frequency performance of the circuit by providing a quick path to ground for any stray radiation from the antenna.Figure 5 – Completed ModelThe completed model shown in Fig. 5 was produced using stereo lithography. After the insert was rigorously cleaned, the ICs and discrete components were inserted into their designated cavities. Conductive silver ink was then dispensed throughout the design to create the circuit interconnections.

Conductive silver ink was then dispensed throughout the design to create the circuit interconnections. After the conductive silver had cured in a convection oven, the antenna and ground plane were coated with the silver ink and again the insert was placed into a convection oven for final curing.Helmet Insert Accelerometer System Demonstration The demonstration included a conformal structure that fits tightly to the interior of a soldier's helmet and has three bolt holes that correspond to holes in the helmet. Fig. 6 illustrates the insert. A wireless-capable RF PIC microcontroller was included that only requires an external loop antenna (seen in the figure as a thick trace with three sides) and matching circuit to be included externally.

The chip includes an internal clock, two general purpose outputs and three analog to digital converter input pins as required for reading the accelerometer – providing an analog voltage proportional to the acceleration of the insert for all three dimensions. A single 3V power pin and ground were also included and a ground plane was painted behind the circuit to improve RF performance. The microcontroller was programmed in assembly and included non- volatile memory in order to store the program with no additional configuration chips required.The software begins by configuring the analog to digital converter and then initiates an endless loop, which repeatedly measures, digitizes, and stores each analog voltage. A 72 bit digital word is then formed consisting of the transmitter serial number (used for device identification at the receiver), function codes, and the three values which acceleration correspond to the three axes.

The then uses Amplitude Shift Keying (ASK) to modulate a 315MHz carrier signal and transmit the 72 bit word along with framing pulses for synchronization.transmitterThe microcontroller for the receiver was also programmed in assembly language. The basic operation of the receiver is program validate transmissions by incoming timing framing pulses, verifying function codes in the transmission, reading acceleration values from the 72 The bit word received. receiver can be configured via software to constantly outputtotheandFigure 6 – Accelerometer System Helmet Insertthe acceleration values to a binary LED display, output acceleration values to the display only if they exceed a programmed threshold, or output the values to an RS232 serial port for use by an external application.ConclusionThis paper has described a novel approach for providing a conformal dielectric substrate, which includes embedded electronics functional with a sensor, microprocessor and antenna.

This novel capability is well suited for bio-medical devices that need to be conformal to the human body and can be customized on a unit level to form fit an individual's anatomy. A three- dimensional off-axis magnetic flux sensor was demonstrated to illustrate the utility of this proposed capability.Future WorkFurther work is required to automate many of the steps in this proposed design process which will convert the output of more traditional electronics printed circuit board CAD. We have had preliminary success making these 2D outputs conformal by using the deformation tool provided by SolidWorks (a simple example is illustrated in Fig. 7) and employing a library of chip and passive sockets that hold chips with press fit and physically separate the pins with 75 micron isolation walls - thus avoiding shorts when dispensing ink to the individual pins and avoiding ink seepage under chip.Figure 7 – Conformal ApproachAcknowledgmentsThe facilities within the W.M.

Keck Center for 3D Innovation (Keck Center) used here contain equipment purchased through Grant Number 11804 from the W.M. Keck Foundation, a faculty STARS Award from the University of Texas System. This material is based in part upon work supported through the Mr. and Mrs. MacIntosh Murchison Chair I in Engineering, and a research contract from the U.S. Army Space and Missile Defense Command and the Homeland Protection Institute to the UTEP Center for Defense Systems Research. The findings andopinions presented in this paper are those of the authors and do not necessarily reflect those of the sponsors of this research.References 1. E. De Nava, D. Muse, S. Castillo, M. Alawneh, M. Navarrete, A.J. Lopes, E. MacDonald, and R.B. Wicker “3D Off-Axis Component Placement and Routing with Solid Freeform Fabrication,” Proceedings from the 19th Annual Solid Freeform Fabrication Symposium, University of Texas at Austin, August 3-5, 2008.2. F. Medina, A.J. Lopes, A.V. Inamdar, R.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceElectropolishing of Ti6Al4V Parts Fabricated by Electron Beam Melting L. Yang1, Yan Wu1, Austin Lassell1, Bin Zhou2 1Department of Industrial Engineering, University of Louisville, Louisville, KY 40292 2Department of Mechanical Engineering, Tsinghua University, Beijing 100084, ChinaAbstract In this study Ti6Al4V samples fabricated via electron beam melting were subjected to electropolishing under various polishing conditions using the re-designed in-house electropolishing device. The surface finish of the processed samples were characterized, and the samples were evaluated via experimentation.

The surface finish of the processed samples were characterized, and the samples were evaluated via experimentation. The results provided preliminary information about the effect of electropolishing on surface crack initiation of the electron beam melted Ti6Al4V parts.thetension-tension fatigue properties ofKeywords: Surface treatment, electropolishing, electron beam melting, Ti6Al4V, fatigueIntroduction Powder bed fusion additive manufacturing (AM) processes have been widely used for direct fabrication of metal structural components that are used in various industries such as aerospace, biomedicine and automobile. It is expected that AM could bring about significant performance improvement and cost reductions for aerospace component manufacturing [1, 2], and various aerospace companies such as GE Aviation have already started to mass produce AM components in structurally non-critical parts [3, 4].

However, use of metal AM components in structurally critical parts is still limited for short-term use and demonstrations [5-7]. This is partly caused by the relatively low fatigue life of AM metal components as results of both internal and surface defects [8, 9]. Surface defects are intrinsic to powder bed fusion AM processes, which is mainly caused by the partial sintering/melting of the powder particles at the surfaces of the parts due to heat dissipation effects as shown in Fig.1. These surface imperfections often become crack initiation sites upon cyclic loading, which significantly impair the fatigue strength of the structures. It has been suggested that surface defects might be one of the major contributors of the low fatigue strength of metal parts fabricated by powder bed fusion AM processes. [8, 10, 11].

It has been suggested that surface defects might be one of the major contributors of the low fatigue strength of metal parts fabricated by powder bed fusion AM processes. [8, 10, 11]. For Ti6Al4V this issue could be further aggravated due to the high notch sensitivity of the material.Compared to laser melting processes, electron beam melting (EBM) processes usually exhibit higher surface roughness due to higher processing temperature, larger powder particle sizes and larger energy beam dimensions. The surface roughness of the EBM parts spread over a broad range but could often exceeds 30µm depending on various factors such as process parameters, material types, powder characteristics and surface orientations [15, 16]. As a result, the typical fatigue endurance of EBM parts are often only about 10% of that of wrought parts when the surface is not treated [9]. In order to achieve satisfactory long-term performance, surface treatment is necessary.1333b.

In order to achieve satisfactory long-term performance, surface treatment is necessary.1333b. Laser melted Ti6Al4V [13] Fig.1 Surface finish of metal powder bed fusion AM partsc. Electron beam melted Ti6Al4V [14]a. Laser melted 316L [12]In previous studies, the authors investigated the feasibility of utilizing electropolishing for the direct treatment of Ti6Al4V parts fabricated by EBM [17, 18]. Through limited amount of experimentation, it was found that the electropolishing process has the potential to significantly improve the surface quality of the EBM Ti6Al4V parts, although the surface polishing effects exhibit rather large variability, and the experimental conditions were rather difficult to repeat accurately. Additional design efforts were dedicated to the improvement of the experimental devices and the setup of experimentation in the attempt to reduce the random errors of the process.

Additional design efforts were dedicated to the improvement of the experimental devices and the setup of experimentation in the attempt to reduce the random errors of the process. In addition, based on the established electropolishing parameters, fatigue testing samples were polished and tested in order to evaluate the effect of surface treatment on the fatigue performance improvement. This paper will present these works in details.Electropolishing surface treatment: Theories Electropolishing process is an electrochemical process in which chemical corrosion of the anode through electrolyte takes place with the assistance of electrical currents. Traditionally electropolishing process is mostly applied on pre-treated surfaces to achieve further improvement of surface finish. Electropolishing effect can be categorized into macromachining and micromachining, which are driven by different mechanisms.

Electropolishing effect can be categorized into macromachining and micromachining, which are driven by different mechanisms. As shown in Fig.2, macromachining, or anodic leveling, is effective in removing macroscopic surface features such as surface waviness, and micromachining, or anodic brightening, is effective in removing smaller surface features (<1µm) and achieving mirror-like surface finish [19]. Both types of electropolishing are largely controlled by the mass transportation processes in the electrolyte, which in turn is often driven by the combined effect of surface diffusion barrier layer formation and surface ionic diffusion.a. Initial profileb. After macromachining Fig.2 Progression of electropolishing [19]c. After micromachining1334During the electropolishing process, the electrolyte regions near the anode surface will accumulate higher levels of ionic reaction products, which potentially serve as diffusion barriers for further reactions.

In addition, anodic passivation caused by the formation of surface metal oxide films could potentially introduce additional diffusion barriers [20]. The diffusion barrier layer inhibits cationic diffusion from the anode and therefore limits the electropolishing rate.The dynamic balance during the electropolishing process determines the thickness and characteristics of the diffusion barrier layer. As shown in Fig.3a, when no fluid flow is present, the profile of the diffusion layer would be hypothetically determined by the profile of the anode surface driven by the diffusion laws. However, due to the presence of electrolyte flow in reality, the actual profile of the surface tend to become thicker towards the recessing features (e.g. valleys) on the surfaces as shown in Fig.3b. As a result, the cationic diffusion in these areas are more inhibited compared to the protruding areas, which consequently leads to gradual leveling of the surface [21].

As a result, the cationic diffusion in these areas are more inhibited compared to the protruding areas, which consequently leads to gradual leveling of the surface [21]. The electrolyte flow can be driven either by diffusion and convection or by external agitation, which makes it possible to adjust the surface diffusion layer profile during the electropolishing process. This unique characteristic makes electropolishing promising in removing surface features at different scales and improving surface finish for AM metal parts.a. Without flow/disturbanceb. Typical, with flowFig.3 Diffusion layer profile in electropolishing [21]Typically the ideal formation of the surface diffusion layers occur within certain process windows. Under a certain temperature and experimental setup, the voltage-current curve of the electropolishing often exhibits the characteristic shown in Fig.4.

Under a certain temperature and experimental setup, the voltage-current curve of the electropolishing often exhibits the characteristic shown in Fig.4. A plateau stage exists at medium range of voltage levels, which correspond to the optimal electropolishing window in which dynamic balance is achieved for the surface diffusion layer [19-21]. Very high voltage potentially results in the breakdown of the surface diffusion layer and over- polishing which is termed pitting, while very low voltage might be inadequate to drive the reactions.

Very high voltage potentially results in the breakdown of the surface diffusion layer and over- polishing which is termed pitting, while very low voltage might be inadequate to drive the reactions. As mass transportation processes are strongly influenced by temperatures, the optimal processing window for electropolishing might alter at different temperatures, which must be carefully monitored.1335Fig.4 Typical electropolishing voltage-current curvePrevious literatures of electropolishing with Ti6Al4V were mostly carried out with highly hazardous water-based electrolytes such as hydrofluoric acid and perchloric acid in order to overcome the significant oxide formation issue as a result of the reactivity of titanium [22-24]. Recently, non-aqueous electrolyte based on ethyl alcohol and isopropyl alcohol was reported to be successfully used for Ti6Al4V electropolishing [25].

Recently, non-aqueous electrolyte based on ethyl alcohol and isopropyl alcohol was reported to be successfully used for Ti6Al4V electropolishing [25]. Due to the lack of water, the oxide film formation could be significantly reduced, which might have contributed to the absence of plateau stage in the voltage-current curve. As a result, a moderate level of voltage was determined as the processing parameter, which generated 6- 8kA/m2 of current density during the electropolishing. In previous studies by the authors, the same electrolyte recipe was followed, and the processing voltage was also determined in a similar manner without the observation of apparent plateau stage in the voltage-current curve [17, 18].Experimental setup Multiple iterations were taken in the design of the electropolishing systems.

While a rather rudimental system as shown in Fig.5a was used during the early feasibility study, it was quickly decided that such setup lacks adequate basic control with the processes. In this setup, the electrodes were held by a plastic base plate that was submerged in the electrolyte, and a stirring bar was used to introduce fluid stirring effect in order to promote electropolishing reaction. Due to the concentric flow pattern introduced by the stirring mechanism, the placement of the samples in the middle of the container was considered to be rather ineffective in ensuring consistent electrolyte refreshing at the anode surfaces. In addition, it was also observed that the plastic substrate deteriorated quickly under the severe processing conditions.

In addition, it was also observed that the plastic substrate deteriorated quickly under the severe processing conditions. Therefore, a re-design was devised as shown in Fig.5b, in which the fixation of the electrodes was realized upside down through a plastic holding plate, and the anode was placed towards the edge of the container, which allows it to align better with the electrolyte flow pattern. The holding plate was designed so that multiple configurations for electrode distance could be selected.1336a. Preliminary systemb. Re-designed systemFig.5 Stirring bar assisted system designThe redesign as shown in Fig.5b was used to study the relationships between various electropolishing parameters including voltage, temperature, electrode spacing and polishing time on the final surface quality of the EBMTi6Al4V samples.

Promising results of surface roughness lower than 100nm Ra was realized, although it was also found that the sample were polished at significantly different rates in different locations as shown in Fig.6a [17]. This was concluded to be caused mostly by the irregular flow pattern of the electrolyte using stirring mechanism. In fact, sufficient rotation-induced bottom stirring introduces vortex flow in the liquid in a beaker, which exhibits axially symmetric flow that circulates from the bottom center towards the bottom edge and then ascend to the upper surface of the beaker before flowing down along the central axis again, as shown in Fig.6b. In addition, due to the exothermal reaction that occurred during the electropolishing process, the electrolyte needed to be constantly cooled by placing the beaker in an ice- water bath, which became insufficient when high voltage was used for the electropolishing.a. Sample [17]b.

Sample [17]b. Stirring induced vortex flow [25]Fig.6 Non-uniform electropolishing via stirring mechanismIn order to further improve the electrolyte flow uniformity, a new circulating flow control system was designed. Fluid pump driven flow control was introduced. In the selection of pump mechanism, rather stringent material restrictions were the electropolishing chemicals. Due to the use of strong oxidizing salts (zinc chloride and aluminum chloride), organic solvent (alcohols), and the existence of electrical voltage, the material used as fluid channel must be both chemically inert and electrically insulating. Asimposed by1337shown in Fig.7a, a rectangular-shaped glass container with an inlet and outlet in opposite sides was custom made and used as the electropolishing container. A Liebig condenser heat exchanger was used as the cooling mechanism, and ice water or water with either controlled temperature was used as the cooling media.

A Liebig condenser heat exchanger was used as the cooling mechanism, and ice water or water with either controlled temperature was used as the cooling media. The circulation of the electrolyte was driven by a peristaltic pump, which allows for isolation of fluid from the pump components in order to avoid corrosion. Silicone fluid tubing was used due to its good chemical resistance to both alcohols and oxidizing salts.a. Container and heat exchangerb. Circulation systemFig.7 Electropolishing device with pump driven circulationIt was soon realized that the cooling efficiency of such setup was still insufficient to cool the electrolyte once the electropolishing process initiated. In order for the Liebig condenser to become efficient, large heat exchange contact areas and longer contact time are desired. However, in the current study, due to the need of higher flow rate for efficient electropolishing, the heat exchange time is rather limited.

However, in the current study, due to the need of higher flow rate for efficient electropolishing, the heat exchange time is rather limited. The rather short heat exchange contact length of the Liebig condenser was determined to be insufficient, and a replacement Graham condenser with helical coil was implemented. The latest version of the electropolishing system is shown in Fig.8.Fig.8 Electropolishing system with Graham condenserElectropolishing experiments In this study the electrolyte was made by following the recipe provided in previous literatures: 1L of electrolyte should consist of: 700mL of ethyl alcohol, 300mL of isopropyl1338alcohol, 60g AlCl3 and 250g ZnCl2 [26]. All the chemicals used for the experiments are anhydrous. The DC voltage-current curves of different electropolishing system setups were evaluated at room temperatures for multiple electrode spacing setting.

The DC voltage-current curves of different electropolishing system setups were evaluated at room temperatures for multiple electrode spacing setting. As shown in Fig.9, no obvious plateau stage was observed, which agrees with the previous study with this type of electrolyte [26], which was attributed to the lack of formation of the oxide films. For the system with pump driven flow circulation, 60-80V was selected to be the operational voltage range, which correspond to about 6-8kA/m2 of current densities for the samples used for this study. This range was considerably higher than the values used in previous studies [26].Fig.9 Voltage-current responses of different electropolishing systemsPreliminary works determined that the electrode spacing is less critical in the polishing process [17]. In addition, smaller electrode spacing might reduce the fluid flow between the two electrodes and therefore affect the electropolishing efficiency.

In addition, smaller electrode spacing might reduce the fluid flow between the two electrodes and therefore affect the electropolishing efficiency. As a result, for the consequent experiments, electrode spacing of approximately 7.5-8mm was used without further modification.A full factorial experimental design was carried out with three input variables: voltage, temperature and polishing time. The levels of each input variables are shown in Table 1. Among the three input variables the temperature is the most difficult one to control. Therefore, during the experiment intermittent polishing was used, which allowed additional time for the electrolyte to cool down.Input variable Voltage (V) Temperature (°F) Polishing time (min) Table 1 Input variables for electropolishing experimental designLevels 60, 80 80, 100 5, 10, 20Samples with flat plate design were fabricated in Arcam EBM S400 system with Ti6Al4V using the system default parameter setting.

All the samples were oriented in such a way that the surfaces to be polished are aligned vertically in the building chamber,1339which results in a rather rough surface finish. The surface roughness of the as-received samples were measured via the Dektak 8 profilometer, and the results showed a roughness of 30-40µm. It was suspected that some of the samples were rougher due to the vacuum issue with the EBM system. Two samples were polished for each electropolishing parameter setup. After electropolishing, the surface roughness of both side of the samples were measured using the same method, and the results are shown in Fig.10. In general the most significant factor appears to be the polishing time since a broad range of polishing time was investigated. Non-uniform polishing was observed despite the improved electrolyte flow control, which was likely due to the non-uniform interaction between electrolyte and the samples at different locations.

Non-uniform polishing was observed despite the improved electrolyte flow control, which was likely due to the non-uniform interaction between electrolyte and the samples at different locations. As a consequence, the front side and the back side of the samples were polished differently. The surface roughness of the samples after polishing for 20 minutes were approximately 5-15µm for the front sides and 0-15µm at the back sides. The combination of voltage and temperature did not appear to have as significant effects, although the 80V-80F and 60V-100F combinations showed slightly better surface roughness consistency. This was intuitive since that both higher voltage and higher temperature would promote the electropolishing rates, while their combination might introduce excessive polishing effects.a. Front sideb.

Front sideb. Back sideFig.10 Surface roughness of electropolished samplesAs one of the benefits of interest of electropolishing is the fatigue performance of the EBM parts, additional experiments were carried out with both 80V-80F and 60V-100F parameter combinations to evaluate the impact of electropolishing on the fatigue endurance of the samples. Fatigue samples as shown in Fig.11 were designed and fabricated using the same EBM parameter setting that was used for the samples for the polishing study. The fabricated samples were also shown in Fig.11. An additional feature was designed to facilitate the gripping of the samples during electropolishing, which could be subsequently removed without affecting the part quality.Fig.11 Fatigue sample1340Multiple samples were polished for both the 80V-80F and 60V-100F parameter setting. In order to maximize the polishing effects, all the samples were polished for 20 minutes.

In order to maximize the polishing effects, all the samples were polished for 20 minutes. Due to the rotationally symmetric feature of the fatigue sample, during the electropolishing process the samples were rotated four times with 90° of rotation angle in the attempt to achieve more uniform polishing around the entire parts. The samples were then machined to create smooth gripping surface. Fig.12 shows some the polished samples after machining. The samples were tested in an Instron Electroplus E10000 fatigue tester under tensile-tensile cyclic loading.Fig.12 Electropolished fatigue samples after machiningBefore the fatigue testing of the electropolished samples, the as-received samples were tested under different loading ratios R in order to establish preliminary ideas with the fatigue performance of the control group. After evaluation it was found that the as- received samples exhibit drastic reduction of fatigue life when the R ratio was between 0.25 and 0.3.

After evaluation it was found that the as- received samples exhibit drastic reduction of fatigue life when the R ratio was between 0.25 and 0.3. In addition, it was found from the fracture surface analysis that for samples tested with R=0.25 the prevailing failure mechanism was surface cracking, while for samples tested with R=0.3 internal defect cracking appeared to be more dominant. As a result, loading ratio of R=0.25 was determined to be used for all the electropolished samples in order to evaluate the effectiveness of the surface treatment process in reducing surface cracking under cyclic loading. The results are shown in Fig.13. Although limited data points were available, the fatigue life of multiple electropolished samples achieved over 1.6 million, which correspond to over 100% improvement compared to the as- received samples (~450000).

It must bet noted that most fatigue testing for the electropolished samples stopped after about 1.6 million due to the time constraints, and only one sample polished by 60V-100F parameter combination was tested to 3.3 million cycles without part failure. Therefore, it could be expected that the beneficial effect of electropolishing surface treatment could be even more significant. On the other hand, the fatigue strength of the electropolished samples appeared to scatter more significantly compared to the as-received samples. Further investigation will be carried out to further identify such effect, however it was speculated that this was partly caused by the non- uniform electropolishing and the formation of new surface defects.1341Fig.13 Fatigue properties of the electropolisehd samplesConclusions In this project multiple electropolishing systems were designed and evaluated.

It was found that electrolyte flow control and temperature control were the two most significant process control variables. Circulating flow driven by peristaltic pump provides protection to the pump and improved flow control with the electrolyte. In addition, a Graham condenser heat exchanger was also implemented in order to improve the cooling of the electropolishing process.Preliminary mechanical testing showed impressive results, with fatigue life increased by over 300% for multiple samples polished at both 80V-80F and 60V-100F processing parameter combinations for 20 minutes.

Although the sample size was relatively limited, the preliminary observations showed promising potential of the electropolishing process in improving the mechanical properties of Ti6Al4V parts fabricated by EBM and other powder bed fusion AM processes.Acknowledgement This work was partially supported by NASA Marshall Space Flight Center Contract #SC- 2019-004 and NASA Kentucky EPSCoR Grant RID-3 #7800003125. The authors would like to acknowledge the support of Rapid Prototyping Center (RPC) and Conn Center of Renewable Energy at University of Louisville.Reference[1] D. Godfrey. Additive manufacturing (AM) and the Honeywell global initiative. Proceedings of the International Society of Air Breathing Engines Conference, Phoenix, AZ, 2015. [2] GE Global Research. http://www.geglobalresearch.com/innovation/3d-printing- creates-new-parts-aircraft-engines. Accessed July 2016. [3] GE Reports. http://www.gereports.com/post/116402870270/the-faa-cleared-the-first- 3d-printed-part-to-fly/.

e Electrostatic Application ofPowder for Selective Laser Sinteringby Lawrence S. Melvin III and Dr. J. I. Beaman University.of Texas. at Austin Department of Mechanical EngineeringAbstractThe electrostatic powder application system is designed to deliver powderto the work area in even well packed layers. It delivers the powder·from a storage container using compressed air, forces the particlestbrougbanelectricfield~ and sprays them onto. a .ground~dplate./.Ajqualitativetex~rimentintO the feasibility and performance of the system is presented. A model of the particle charging mechanism,andtheParticle charge draining mechanismis also<presented. Theexperinlentfmds that a 20 mill layerof packedpowderisrepeatedly achievable, and that the system has powder clogging problems.IntroductionThe electrostatic powder application s~stemis intendedto replace the current roller feed mechanism inthe··.selective< lasersinterin~process ..

Theelectrosta~csystem.should help to create~tt~rpackedpowderlayers,.·redu"ethework.•spacer~uiretnents, .a.nd·all0wthe.poW(jer and to bestoredinjadifferenten.virontnentfromthe\Vor.k~paceJj>riortothe.~pplication sintering process. These•.factors .sh()uldhelptoincre~sethe9uality.ofthesinteredpartbY reducing the number of air pockets in the work piece, and by helping to.keep.applied layers even.TheoryA theoretical model of the electrostaticpowder application.process has been derived to determine if .the desired effects are .being produced atthe particlelevel. •. Several assumptions have .·beenmadeto produce a simple closedform••mod~lof~lttavelin.~cbargetiparticle.First, has been assumedthatthe particlesa.re spherical. Second, the modeled<particle\Villbe in the center of the flow.

Second, the modeled<particle\Villbe in the center of the flow. Finally,the groundedworkpl~teisasst11lledtobea.~rfectcol1~uctor,and any.l?reviouslyaJj>Jj>lie<i.po\Vderla.yers.are.charge.netltral.t.'I'he!"esultingilpQdelwill lJe able to describe •how the<particles.are influenced, but the magnitudes ofthe results should be considered suspicious and only used to compare particles and Illaterials, not for prediction of physical behavior.The capacitance of the material particle iSl1eededtodeterrnirt~the charge that is held by the particle, to resolve the force on the particle, andto.deterrrri.nethetime needed fortheqhargeto drain from the particle after the/particle has been applied. A modelis derived from the basic 1vIRXwell'sec.ttla.tiQns:41t£oRQ= V41t£oR171it(1),(2).These equations yield the capacitance of, and the charge on a given particle.The electrostatic force that is generated by the charged particle is calculated using the methodof images.

The method finds the radial and tangential components of the electric field to produce the field in the normal direction:.2QD -................ i........ En = Ercos9 .. E.esin9>= ••••• ••• 3 En41teofThis expression for the electric field can be used to calculate the force that is acting upon the charged material particle, refered to here by its charge Q:Theradiusofthesystelnisknown,r:=12D,and~a.nbesubstitut<edinto equation. 3, the result of which can be substituted into equation 4 to produce an expression of the electrostatic force upon the particle:A particle that is traveling between the spray head and the work plate is influenced by the electrostatic force, the drag force of the fluid, gravity, and the force of the fluid stream. The gravity force is simply the mass multiplied by the gravitational constant, while the electrostatic force was derived previously. The fluid stream force occurs in the applicator system, not the area between the feed head and the work plate.

The fluid stream force occurs in the applicator system, not the area between the feed head and the work plate. Assuming that the particle velocity is the same as the fluid velocity at the time of exit from the spray nozzle, the fluid imparts an initial momentum upon the particle equal to the product of the fluid velocity and the particle mass. All particle particle interaction forces are assumed to have a resultant vector of zero.The particle traveling through the air is also acted upon by a fluid drag force from the air around it, which is proportional to the difference between the particle velocity, and the velocity of the air around it.12 2 Fd=2 pACrJ.AV)All ofthe forces can then be summed to produce a model ofthe influences upon the particle:The next step is to createan.expressionforthevelocity.frQlnthe force equation.

To createthe velocity.expression, first,the.!Qrceequation will be divid<edby the mass ofthe particle, thentheconstallts in the equationvvill.belumpedtQgether into>three constants a, b, and c. This yields·a simple expression for the acceleration·ofthe particle:172(3).(4).(5).(6).(7)..... b .2 a(x) =2 + c + dxxThis equation has two nonlineari~e~jnitduetothecomponeJ.1~98ntributed by the drag force, and the component due to the el~,trostatic forcy,thuselirninatingthepossibility of a closed fonnsolution to this problem. However, th~drag force does not~ecessarily make a significant contribution to the acceleration of the particle, because itdepel1dsupon on the difference between the velocity of them it istrave~gwithandthevylocity of the particle itself. This difference is notgoingto be large bec~u~etheparticleonl~travels a smalldistance between the spray head and the work plate, not allowingtl1Y(diff~rencetoincreaserapidly.

This difference is notgoingto be large bec~u~etheparticleonl~travels a smalldistance between the spray head and the work plate, not allowingtl1Y(diff~rencetoincreaserapidly. The drag force tenncan then be dropped from the system model without10sing unacceptable generality or detail:b .... a(x)=-+c 2 xThe velocity of the system can be found using simple dynamics. The velocity as a functionofparticle position will be integrated up from the'acceleration expression:..

a(x)=-+c 2 xThe velocity of the system can be found using simple dynamics. The velocity as a functionofparticle position will be integrated up from the'acceleration expression:.. xdvdv =dx dt ="dXdv dxf·· '~:2+C)dx=(VdVXoThis expression is now integrated to obtain an expressionfor the velocity of the particle:b xxo2 v =vo- 2(x - xO)I- + c]2The initial velocity is the velocity of the charged particle at the exit of the spray nozzle, and initial position of the particle is the vertical displacement at the exit of the spray nozzle.Once the particle has landed upon the work surface, the charge will be drained to ground.This system can be thought of as a simple current driven RC circuit.

The resistance and capacitance of the layers is needed to describe the charge dissipation, realizing that different values of these terms, are present due to the state of sintering and the presence of boundary conditions..An important reason for this theory is the hypotheses that a charged particle in flight willtend to fill in a valley. The model that is derived here assumes that the valley is a two dimensional cup, and that the charge on the walls of the valley can be approximated by four point charges with charge of the same sign as the particle's. This model is an excellent approximation because the walls of the cup consist of discrete charged particles.

This model is an excellent approximation because the walls of the cup consist of discrete charged particles. This model assumes that all of the particles, those in the valley and the travelling particle, are being applied during the same feed head stroke.173(8).(9).(10).(11).(12).Particle>approachinga ValleyFigure 1ThecllargeQntheparticlesinthecupwatlwill.•bedifferentduetothe difference in the timebetweenimpactofthe particles, and the difference in the powder thick11essbelow them. An expression of their charge is:t 1 ql =q4 =Qoe- RC t 2 q2= q3 = Qoe- RCwhere t2>t1. The forceactingiupon thetravelling particle by the four charged particles in the wall of the cup is:200s9 n1-4xeoXn Fn~ -qn~~ine.j 4XEoXn.. F nQX=·qnQA.An example using numerical values in equation 7 is worked using values for nylon 6 and alumina. Both of these materials have been sintered.

F nQX=·qnQA.An example using numerical values in equation 7 is worked using values for nylon 6 and alumina. Both of these materials have been sintered. The values used in the equation are from reference material containing information about these two materials. These numbers can be substituted into simplified equation 7, and produce a numerical representation of the interaction between the plate and the particle:Fn=-10 1.80 xlO A. 4 69. 2lO-8 A JI>Fa=-4.55 x 10 I>210~ 1 11 J-.x10-7A,.

The initial velocity is one that was typically used in the experiment. The final value of the displacement was 0.1 mm. The results of the velocity simulation are graphed in figure 2.The velocity of both the alumina and the nylon particle theoretically approach infinity. The acceleration of the particles at a distance of 1 mm from the work surface is 39 g for the nylon particle, and 42 g for the alumina particle. This large increase in the acceleration of the particles begins at a distance of 7 mm from the surface of the work plate, a distance of approximately 35 particle diameters from the work surface.The following example illustrates the forces acting upon the charged particle when the particle approaches a valley of charged particles.

The values of resistance and capacitance for the charged layers of material were chosen based on the layer thickness of the real particle layers, and known sizes of capacitors and resistors making a reasonable approximation of the contribution of the air gaps between particles to the resistance and capacitance of the material. The positions of the charged particles were chosen by placing them in a hemisphere of radius 0.005 m, and placing the travelling particle at the point 1 mm to left of the center, and 1 mm above the semicircle. It is assumed that the particle has momentum only in the negative Y direction when it is at this point. Two values of time are used to compute the charge on the particles in the valley, and the values are based upon the particle velocity.These numbers were put into equations 29 and 30 and summed over the particles. For thenylon system, the numbers are:~~.

For thenylon system, the numbers are:~~. F n = 1.82 x 10 1 + 4.69 x 10 IN an =3,807.5 i + 10,376.5 j ~ s.~.~....AAmThe alumina numbers are much smaller:Fa = 1.23 X 10 i + O.90x 10 jN _7 A....AAmaa = 10.9 i+ 7.96 j ~ s175(33),(34).(35),(36).nylon, which has a larger electrical time constant, will be influenced muchExperimentAn electrostatic powder application system was constructed and tested to determine if the desired physical phenomena were present. The apparatus consisted of a powder canister which holds and delivers powder similar to an aerosol can, flow control devices that regulate compressed air flow into the powder canister and the air flow that carries the air from the powder canister to which is mounted on a linear actuator rated at 20 ips. The feed head turns the flow from vertical to horizontal, forces it through a spray nozzle, and electrostatically charges the powder.

The feed head turns the flow from vertical to horizontal, forces it through a spray nozzle, and electrostatically charges the powder. The powder is charged in the spray nozzle with two 13.5 kV, 0.31 rnA power sources, one delivering positive charge the other negative. A set of valves were inserted in the line between the feed head and prevent Thefeed head. The powder leaves the canister and enters the feed head,powder canister to bleed most of the powder flow off in an effort tothe C1'l'C11b::l>1I"Ir"l was performed with four variables, the feed and carry pressure, and thefeed and carry flow rate. The other four variables, feed head speed (16 ips), feed head height and material (PVC), were all kept constant. The sintering was simulated (7 in.), spray using a experiments were conducted on a qualitative basis because the clogging problem experiment was The data applied and visible. When only a few surface. When fewer.

When only a few surface. When fewer. The two layerstorch.efforts to produce a consistent powder layer, while a simulated roller conducted for comparison.experiment was interesting. When the frrst layer of powder was it formed several holes in the surface, through which the work plate was powder was applied, the holes were gone. After sinteringlayer the work plate to be exposed, and there were also a few valleys in the simulated roller work pieces, the holes were shallower, and with ranged from 12 to 16 mills for a single layer and 20 to 28 mills for of two or three mills on each work piece.a An error wasduring the experiments that lead to the most interesting discovery. Onelayer of powder was applied with all of the flow reduction valves off, thus letting a large amount of powder flow out amount of parts fabricated of thethe spray nozzle. Unlike the previous experiments, a largewas floatingthe air around the spray head after the application.

Unlike the previous experiments, a largewas floatingthe air around the spray head after the application. Whenway were measured, they were consistently 24 to 26 mills thick, regardlessa thickness variance of one mill.Clogging waseasily However,closing allsystem each time the powder flow was cut off. The clog was the flow reduction valves and letting the clog blowout.phenomenanot been eliminated from the process.Analysis~e,.,eni:U. ......,u'...........M.U'''''' cJ1ar~:tCU~t1sttcs can be seen in the experiment and the theoretical model of shows that a charged particle will tend to settle in the center of a valleythe process. of like charged particles, thereby filling the valley.· The experiment also demonstrated that a charged particle The layerstend to with theholes or valleys in a work surface.reduction valves off are very informative.

It appears that work surface and builds layer thickness until enougl1 charge is built additional charged powder particles which are attempting to come tothe powder up that it starts rest upon the because this thickness is constant for a given material and charge.onprocess provides excellent filling of the hills and valleysThe clogging phenomena and the inability to produce layers that are as thin as are needed for to the need for a redesigned system. The current systemselective laser sintermg (5 mills) provides information about how the electrostatic process works, but it is not capable of the accurateperfonnance needed in the selective laser sintering process.176SuggestionsThere are severill ways that the system may beredesigned to control the thickness and eliminate the clogging problem.• The clllTentpneumatic SySteIllmay be modified to allow the powder to beforced into the carry flowbytlienaturalvac~umofthe flow over the powder orifice.

Leaving the systernwith an open loopcontroLprocess does notprovide enQugh controlto keep a rigidlayer thickness. The thickness could be controlled using a feedbackmeasur-ings~stem~particlechargecontrol~or a simplemecbanicalscraperorroller.· ·Theifeed·backmeasurirtgsystem would>beaiproc7ss of applying.·and.sintering a layer~rmdingan.dfillingI>roblemspots•with the m7asurin.gsystem~ and thenapplyinganewlayerandsinteringit.•••·Thepfll'ticlechargecontroLwouldus7the patticlelayerfepulsionphenomenato controLthe thickness ofthe layers.•• The scraperis the simplest method~jtsimply removes an applied excess ofpowderto keep a consta.fit layer thickness.ConclusionsThe eleetrostaticpowder application system helped the particles find and ftil valleys in the workpiece~andbelpedwith powder packing. The clllTentsystemis notable to provide thin enoughlayers and sUifersfroIll problems with clogging and therefore~. itl1eeds to be redesigned.

ectrostatic Stabilisation of Drop on Demand Bio‐Ink through the Cationic Encapsulation of CellsMatthew Benning 12* Kenny Dalgarno 11 School of Mechanical and Systems Engineering, Newcastle University, UK 2 EPSRC Centre for Innovative Manufacturing in Medical Devices, Newcastle University, UKMatthew.Benning@NCL.AC.UKAbstract The ability to formulate bioprinting inks in which suspensions of cells and other biological materials can be maintained, without affecting biological response, is crucial in producing robust printing strategies for tissue fabrication. A piezoelectrically actuated drop‐on‐demand printing system has been used to deposit electrostatically stabilised cells from a human osteosarcoma cell line (U2OS). Experiments were intended to investigate the effectiveness of a polyelectrolyte cell encapsulant to maintain cell dispersion within a bio ink.

Experiments were intended to investigate the effectiveness of a polyelectrolyte cell encapsulant to maintain cell dispersion within a bio ink. Cells were coated with a number of thicknesses of a Cationic Poly‐l‐lysine (PLL) encapsulant and their ability to release studied over 7 days, with the thinner coatings proving to be more favourable. Printing of both coated and uncoated cells indicated the dispersion and printability of coated cells was significantly better than that of uncoated cells. Preliminary results suggest that electrostatic stabilisation of bio inks could provide a solution to cell aggregation, increasing viable printing time and decreasing poor yields due to orifice obstruction.Introduction Tissue engineering and regenerative medicine aim to promote the regeneration of tissues following damage or disease by the provision of scaffold/support materials incorporating appropriate cells and biological molecules.

In order to deliver these aims, innovative delivery mechanisms and processes need to be developed. Additive layer manufacturing (ALM) is the collective term that describes production techniques which allow the construction of complex geometries through the assembly of 2.5D layers. The use of digitally controlled positioning systems to precisely deposit biological materials within scaffolds has been the topic of research for two decades. To this end, Bio plotting, the process of low temperature extrusion, has allowed researchers to extrude cell rich hydrogels in to predetermined 3D geomiteries1. While the low resolution grading of a number of materials is possible within bio plotting type systems, it adds significant complexity to the manufacturing system and results in lower productivity.

While the low resolution grading of a number of materials is possible within bio plotting type systems, it adds significant complexity to the manufacturing system and results in lower productivity. Inkjet printing on the other hand is widely accepted within the graphics industry as a high throughput printing system capable of grading multiple materials at high resolution with little or no impact to productivity. Cells to be inkjet printed are suspended within a carrier fluid and as a mixture are referred to as a bio ink. The rheological properties of these inks and their impact on printing have been well documented2 but the specific challenges associated with the flocculation of cell within bioinks has not been addressed. Brownian motion, a continuous time stochastic process, describes the random motion of particles through a liquid due to their continuous bombardment from the liquid’s constitutes molecules3.

Brownian motion, a continuous time stochastic process, describes the random motion of particles through a liquid due to their continuous bombardment from the liquid’s constitutes molecules3. Electrostatic stabilisation using polyelectrolyte coatings is a common method of stabilising colloidal pigment suspensions. Many of these coatings however, are cytotoxic and thus inappropriate for use within a bio‐ink. A cationic polymer, Poly‐L‐Lyseen (PLL) is known to be both cell compatible4 and an electrostatic stabiliser5. It was therefore hypothesised that the encapsulation of cells within PLL would aide dispersion, and increase the printability of a bio‐ink.2.

It was therefore hypothesised that the encapsulation of cells within PLL would aide dispersion, and increase the printability of a bio‐ink.2. Materials and Experimental procedure2.1 Cells and Culture Conditions U‐2 OS (human derived osteosarcoma cell line) were obtained from the ATCC and cultured in 75cm2 flasks (Coring‐Costar, CLS3375) in low glucose Dulbecco’s Modified Eagle’s Medium (DMEM) (Sigma‐ Aldrich, D5546) supplemented with 10% (v/v) fetal bovine serum (FBS; Sigma Aldrich, F9665), 5mM Glutamax (Invitrogen‐Gibco, 35050) and 1% (v/v) penicillin/streptomycin (Invitrogen‐Gibco, 15140). All cultures were maintained at 37˚C in a 5% CO2 humidified incubator. Cells were detached every 3 days after they had reached 75‐80% confluence. The media was aspirated from the confluent monolayer and washed with warm sterile 1X PBS.

Cells were detached every 3 days after they had reached 75‐80% confluence. The media was aspirated from the confluent monolayer and washed with warm sterile 1X PBS. The monolayer was incubated with 0.25% w/v trypsin and 0.53 mMol/L EDTA (Sigma‐Aldrich, 25300‐054) for 2 minutes and followed by inactivation of the trypsin with growth medium. The cell suspension was centrifuged at 200g for 5 minutes. The pellet was resuspended in growth medium and cells were reseeded at 1:3 ratio2.2 Cell Coating and Bio‐Ink Formulation For encapsulation, U‐2 OS were suspended in Hank’s Balanced Salt Solution (HBSS) without phenol red in a concentration of 2x106 cells/ml. PLL‐FTIC (fluorescein isothiocyanate) solution (1 mg/ml) was prepared in Hank’s Balanced Salt Solution (HBSS) and filtered using 0.22 μm Whatmann sterile filter. Cationic PLL (Sigma) was first added in three different concentrations: a) 200 μg, b) 400 μg and c) 2000 μg of PLL‐FTIC to a 1 mL suspension of U‐2 OS of 2 million count.

Cationic PLL (Sigma) was first added in three different concentrations: a) 200 μg, b) 400 μg and c) 2000 μg of PLL‐FTIC to a 1 mL suspension of U‐2 OS of 2 million count. Cells were incubated with the PLL during 15‐minute at 37 °C in a 5% CO2, 95% humidified air environment, followed by three washings steps with HBSS at 37 °C using centrifugation at 250g for 5 min to remove any excess or unadsorbed polyelectrolyte. Encapsulated cells were re‐suspended in 1 ml of filtered HBSS media to form a bio‐ink. The uncoated cells were detached from culture as indicated in the previous section and resuspended in filtered HBSS at a ratio of 2x106 cells/ml to form a second bio‐ink.2.3 Printing Experiments Two channels of a Jetlab 4 (Microfab Inc., USA) single orifice piezoelectric printer were used to perform the printing experiments. Droplet ejection was achieved through the actuation of the piezoelectric material within the inkjet head, affected by the application of a potential difference.

Droplet ejection was achieved through the actuation of the piezoelectric material within the inkjet head, affected by the application of a potential difference. The impulse was applied in a simple trapezoidal wave form with a ramp up, dwell, ramp down, amplitude and drop frequency of 3µs, 17µs, 5µs, 60v and 5 kHz respectively. Droplet formation was confirmed prior to printing using a LED stroboscope which was phase shifted from a base frequency matching a factor of the print head frequency. The phase shift allows the tracking of a set of droplets over a given displacement. Aphelion (Microfab) image analysis software was used to calculate the mean velocity, diameter and therefore the volume of the ejected droplets. Following formulation of the two bio‐inks, one containing PLL Coated and one containing Non‐coated U2OS cells, and after a 15min transport from cell culture to the printing laboratory two 10µl sample droplets were taken directly from the micro tubes prior to printing.

The bio‐inks were printed from a 60µm diameter orifice, onto a transparent poly(methyl methacrylate) (PMMA) substrate, from a height of 1mm. 10 Deposits consisting of 50 droplets were deposited in a 1x10 array for each Bio‐ink at 7 time points from 0 to 60mins in ten minute increments. Following deposition of the sample and at each 10min time point, the deposits were viewed through an inverted microscope (1OJ3, Olympus, Japan) and the number of cells recorded.3.0 Results:3.1 Cell Encapsulation Verification of the cell coating was obtained by florescence microscopy where the FITC marker, present within the PLL shell, fluoresced green in the presents of the microscopes ultraviolet radiation source. Figure 1, provides a view of the shell structure at two days post‐encapsulation, showing the adsorption of PLL onto the cell membrane and the formation of an encapsulating shell around the cells.

Figure 1, provides a view of the shell structure at two days post‐encapsulation, showing the adsorption of PLL onto the cell membrane and the formation of an encapsulating shell around the cells. The comparison of fluorescence images revealed a decrease of the coated cell diameter with an increase in PLL concentration during the coating process.Figure 1. Fluorescence microscopy of encapsulated U2‐OS cells at different concentrations of PLL: A) 200 μg/ml, B) 400 μg/ml and C) 2000 μg/ml (per volume of cell suspension), after incubation during 2 days. Scale bar: 25 μm.3.2 Printing Prior to printing the 10µm sample taken from the micro tube reservoirs were viewed under a microscope to visually assess cell distribution. Figure 2, show the comparative cell distribution of the two bio‐inks containing coated and non‐coated cells.

Figure 2, show the comparative cell distribution of the two bio‐inks containing coated and non‐coated cells. The bio‐ink containing the non‐coated cells showed significant cell agglomeration, whereas the coated cells are well dispersed throughout the imaged droplet.Figure 2. Optical microscope comparison of U2‐OS cells in HBSS bioink suspension prior to printing experiments, a) non‐ coated cells, b) cells with a cationic Poly‐l‐lyseen coatingFigure 3, shows an example of droplet formation and one in a series of images from which drop diameter and velocity can be calculate. The voltage/time impulse variables for the bio inks were empirically derived to optimize droplet formation.Figure 3. An optical micrograph taken with a stroboscope demonstrating optimised droplet formation.The bio‐inks were printed in 2 x 10 arrays with each point consisting of 50 droplets. The number of cells at each point was subsequently counted and the numbers recorded.

The number of cells at each point was subsequently counted and the numbers recorded. Figure 4, shows the average number of cells over 10 points at each time point. The non‐coated cells ceased printing after 10mins, and after showing initially very high cell per drop ratios. The coated cells demonstrated a relatively stable cell per drop ratio over the full printing period.Figure 4. A plot showing the average number of cells per 50 drop sample over a 60min time frame.4. Discussion The cell encapsulation process showed that it was possible to effectively coat cells in PLL. The images of the coating, obtained by florescence microscopy, revealed a decrease of the coated cell diameter with an increase in PLL concentration during the coating process. This phenomenon could be attributed to an initial cytotoxic effect caused by the high polymer concentration and coating thickness.

This phenomenon could be attributed to an initial cytotoxic effect caused by the high polymer concentration and coating thickness. This increase could then be preventing the two way permeation of vital metabolic chemicals such as nutrients and waist. The pre‐print contents of the bio‐ink cartridges showed a significant difference between coated and non‐coated cells, with non‐coated showing significant flocculation. As flocculants were not visible during bio ink formulation it is believed that aggregation occurred in the 15min from formulation in the cell culture facility to arrival at the printing laboratory. Conversely, the coated cells showed good distribution through the drop and no evidence of flocculation. Uncoated cells were shown to block the orifice and stop printing after 10mins. The orifice blockage showed that a non‐coated cell suspension would be unsuitable for prolonged printing.

Uncoated cells were shown to block the orifice and stop printing after 10mins. The orifice blockage showed that a non‐coated cell suspension would be unsuitable for prolonged printing. Conversely the coated cells showed a regular deposition of approximately 1 in 25 drops leading to the deposition of a cell. Although the cytotoxicity of the coating was not measured in this work, it is apparent that the U‐2 OS cells coated in the lower concentration PLL survived, broke free and went on to proliferate after 7 days.5. ConclusionIt is possible to coat cells in a Cationic Polymer for use as an electrostatic stabiliser in bio‐ inks.Cells coated in a concentration of 200 μg/ml PLL will break free and proliferate after 7 days.  The stability of bio‐ink cell dispersions can be greatly enhanced by the use an electrostatic stabiliser, and will ultimately aid predictable printing.6. References[1] Kim G, Lee H, Kim Y, Kim S.

EVATED REGION AREA MEASUREMENT FOR QUANTITATIVE ANALYSIS OF LASER BEAM MELTING PROCESS STABILITYJ. zur Jacobsmühlen1, S. Kleszczynski2, G. Witt2, and D. Merhof11Institute of Imaging & Computer Vision, RWTH Aachen University, Germany 2Institute of Product Engineering, University of Duisburg-Essen, GermanyAbstractLaser beam melting (LBM) processes enable layer-based production of geometrically complex metallic parts with very good mechanical properties for Rapid Manufacturing. Col- lisions between powder coating mechanism and elevated part regions pose a major risk to process stability, which is crucial for industrial application. Minimizing elevated region area usually involves parameter tuning in a trial-and-error approach, as the process outcome is the only measure of stability.

Minimizing elevated region area usually involves parameter tuning in a trial-and-error approach, as the process outcome is the only measure of stability. One published approach to quantifying elevated region area utilizes an imaging system, which acquires layer images of the powder bed after powder deposition and detects elevated regions using image analysis. We extend the image-based analysis to each part region, create quantitative visualizations of elevated region area for quick assessment/comparison and compute a ﬁgure of merit. In experimental build jobs with overhanging structures and diﬀerent support junction parameters we gain insight into problematic part regions, which can be used as feedback in job design. The presented method helps to improve LBM process stability, which is strongly linked to process eﬃciency.IntroductionIn laser beam melting processes, parts are built in a powder bed by selectively melting the current powder layer according to the part geometry.

After lowering the build platform a recoater blade or roller moves powder from the powder container onto the build plat- form (Figure 1a) before the laser melts the next layer. All steps are repeated to build the entire part.The possibility to produce complex and individual parts in a tool-less manufacturing process on the basis of CAD data was identiﬁed as the main driver of innovation in [1]. For industrial applications laser beam melting (LBM) is of special interest, due to the pos- sibility to produce complex metal components with suitable mechanical properties. First LBM production lines are already established in the sector of medical technologies [2] and aerospace [3].A problematic eﬀect during LBM build jobs is the buildup of elevations, which stand out from the part layer (Figure 1b).

As the powder layer is very thin (20 to 100 µm) these elevations are not covered by metal powder and may collide with the recoater blade causingLaserRecoater bladeBuild platformPowder containerLEDlightMatte reflectorxzyLaser windowxzRecoater bladePowderPartxy(a)(b)(c)Figure 1: (a) The process chamber with the build platform and the recoater blade, (b) ele- vated part regions (highlighted) cause collisions between recoater blade and part, damaging either one or even both. (c) Example of acquired layer image.damage to either one or even both. Severe elevations may cause process breakdowns by blocking the recoater blade, or destruction of thin parts, which are disrupted by the passing blade.Consequently, the goal during the LBM job design phase is to minimize elevations in order to maximize process stability, which is required for reliable part production.

Process monitoring and quantitative evaluation of elevations is needed to achieve this goal.Several approaches to stability evaluation of LBM processes have been published. Reinarz and Witt [4] measured accelerations caused by contacts between blade and part by means of an accelerometer on the recoater blade. They deﬁned thresholds for acceleration and interrupt the process when these are exceeded.Krauss and Zäh [5] investigated the process reliability by building thin wall structures at diﬀerent angles with respect to the coating blade and manually analyzing an image of the powder bed after deposition of a new powder layer. They observed collisions between part and recoater blade, which lead to mechanical distortions for thin (250 µm) wall structures. Kruth et al. [6] developed an in-axis imaging system for melt pool monitoring and process control, which measures the size of the melt pool area.

Kruth et al. [6] developed an in-axis imaging system for melt pool monitoring and process control, which measures the size of the melt pool area. For overhanging structures the melt pool size increases due to the lower conductivity of metallic powder materials. This leads to melt process instabilities e.g. balling and overheating and consequently shape deviations and elevated part regions, which endanger process quality and stability. The developed system is able to detect changes of the melt pool size and adapts the laser power to keep it constant. This system is extended by mapping the acquired images to layer images and detecting process failures in [7]. The analysis of the mapped melt pool signal is applied to the layer before and the overhang layer itself to measure overheating for varying numbers of support structures.We have presented an approach for the inspection and assessment of process stability in [8].

Here, an imaging system for acquisition of powder bed layer images [9] is used to evaluate the processing behavior for the support transition and overhanging structures by measuring total elevated region area and mean area of connected regions. The image- based measurements were validated against one-dimensional accelerometer measurements and showed good correlation.Compared to the one-dimensional accelerometer measurements, two-dimensional mea- surements from powder bed layer images provide spatial information about elevations, whichhelps to identify the probable cause, e.g. a critical geometrical feature.

a critical geometrical feature. In this publication we extend the analysis of elevations detected in powder bed layer images by using imported reference part geometries and creating quantitative visualizations of elevations for all layers of a build job.MethodWe use an externally mounted monochrome CCD camera (29 Mpixels) with a tilt and shift lens to acquire layer images Iz of the powder bed before laser exposure. The image resolution ranges from 20 to 30 µm/pixel depending on the ﬁeld of view (approx. 150 mm × 110 mm). Indirect lighting of the powder bed is provided by an LED line light at the machine front and a matte reﬂector at the machine back ([9], Figure 1a). It helps to reduce glare and reﬂections on the metallic part regions.Image acquisition is triggered by the limit switches of the recoater blade to obtain layer images of the powder bed.

It helps to reduce glare and reﬂections on the metallic part regions.Image acquisition is triggered by the limit switches of the recoater blade to obtain layer images of the powder bed. There is one powder bed image Pz ∈ ZM ×N for every layer in the build job, the index indicates the dependency on the layer’s z position.The detection method from [9] is based on the assumption that most pixels inside a pow- der bed image Pz show powder. We identify the image intensity as Iz(x, y) with (x, y) ∈ Pz and compute the intensity meanIz =1M NXx, y∈PzIz(x, y),which represents the powder. Then, elevations can be detected as outliers using a detection thresholdTz = 3 · σ, where σ is the standard deviation of all pixel intensities Iz(x, y) and the factor was determined experimentally.

Then, elevations can be detected as outliers using a detection thresholdTz = 3 · σ, where σ is the standard deviation of all pixel intensities Iz(x, y) and the factor was determined experimentally. Tz is computed for every layer and the detection result is obtained as a binary mask imageDz(x, y) = 0 Iz(x, y) ≤ Tz 1 Tz < Iz(x, y),The stride in z direction depends on the layer height, which is a process parameter and may range from 20 to 100 µm. Typical layer heights in our jobs are 40 µm for support structures and 20 µm for part layers.As the recoater blade moves in x direction, large elevated areas at a single position x0 are critical for the process stability as they are hit at the same time. Since we want to quantify the process stability we extract the elevated area Az(x) at every x position inside every layer image from the detection mask Dz asAz(x) = Apixel · XDz(x, y),0≤y<Mwhere Apixel is the area of one pixel in mm2.

Figure 2 illustrates the measurement of elevated region area from powder bed images. A threshold for critical elevated area per x position was deﬁned as Acritical = 0.1 mm2 based on accelerometer measurements [8].(1)(2)(3)(4)input powder Figure 2: Example of elevated region segmentation and analysis. Top left: bed layer image with visible elevated regions and highlighted region (scaled from 29 Mpixel). Top right: close-up of elevations of a single part with support structures. Bottom left: segmentation result based on method from [9] note that regions which are smaller than the laser focus diameter are excluded. Bottom right: integration of elevated regions to obtain Az(x) as a measure of elevation severity at every position x0.Overhanging regions in general and the down-skin in particular are the sections of the build job with the most elevations. Here, we are not only interested in the accumulated area at every position xi but also in the region statistics, i.e.

Here, we are not only interested in the accumulated area at every position xi but also in the region statistics, i.e. the spatial distribution. We analyze regions inside Dz using connected component analysis and compute their area AR,z,i, and the total number of elevated regions NR. For connected component analysis the implementation from [10] is used.The described measurements can be carried out on entire layers, part regions (deﬁned by their bounding boxes) and parts (deﬁned by their geometry in each layer), see Figure 3. We import the part geometry from Common Layer Interface (CLI) ﬁles and store them along with the acquired layer images using the Hierarchical Data Format (HDF5). The analysis of part regions enables the separate analysis of multiple parts in a build job e.g.

The analysis of part regions enables the separate analysis of multiple parts in a build job e.g. to compare process stability of identical parts built in a grid with varied process parameters.For parts with support structures in part layers the analysis of part regions does notFigure 3: Regions for elevation detection analysis (enclosed by dashed lines, elevations are shown in red). Left: entire layer, all elevations are taken into account, center: part regions deﬁned by part bounding boxes, right: part area only, all elevations outside of part area are ignored.diﬀerentiate between part and support structures which is undesirable e.g. if we want to analyze the inﬂuence of diﬀerent support structure junction parameters.

if we want to analyze the inﬂuence of diﬀerent support structure junction parameters. Here, the analysis of parts can be used to analyze these areas separately (Figure 3, right).The availability of part geometries enables computing the relative elevated part area:Arel(x) = Az(x)  Apart(x) 0 if Apart(x) > 0 else,, x ∈ [xpart,min, xpart,max]which can be averaged over the entire part region [xpart,min, xpart,max] and all L layers of a build job to obtain a single value for part comparison:Arel,avg =1LXxArel(x)Quantitative visualization of elevated area for an entire part and all layers of a build job is complicated in a 3D visualization as the results of multiple layers would overlap, making it hard to see e.g. regions with increased elevated area.

regions with increased elevated area. Instead, we propose to combine all elevated regions at speciﬁc positions xi by integrating over yA(x, z) = XDz(x, y),ywhich yields cumulative vertical cross sections (x/z-plots) with color-coded elevated area. These plots enable identiﬁcation of critical layers and localization of critical geometrical part features (Figure 5), and quick comparison of process stability for identical parts built with diﬀerent parameters (Figure 7). A similar visualization was used in [11] for thermograms to identify hot spots.For manual qualitative inspection of detected elevations of an entire build job a 3D vi- sualization can be useful.

A similar visualization was used in [11] for thermograms to identify hot spots.For manual qualitative inspection of detected elevations of an entire build job a 3D vi- sualization can be useful. We use ImageVis3D [12, 13] to visualize exported elevation layer image stacks (Figure 4).(5)(6)(7)Figure 4: 3D visualization of elevations created from image stack using ImageVis3D [13]Experimental ResultsFor build jobs with critical geometrical features the process design has to maximize the process stability. Overhanging geometries for instance are geometrical features with an increased risk of elevations. Using layer image analysis the critical overhang angle and the location of starting elevations can be identiﬁed and counter-measures can be taken, e.g. rotating the part to align with the movement of the powder blade. The optimum rotation angle has to be identiﬁed in experimental builds, which yield only a binary result: success or breakdown.

rotating the part to align with the movement of the powder blade. The optimum rotation angle has to be identiﬁed in experimental builds, which yield only a binary result: success or breakdown. Elevation analysis provides additional insight into the development of elevations during the build job and helps ﬁnding the transition from stable build to unstable build more precisely.To evaluate our analysis methods we build two jobs on an EOSINT M 270 LBM systemby EOS, Germany, using Hastelloy X:job A: 16 identical parts with overhanging geometries (Figure 5a), andjob B: 25 cubes,for which the support structures (number and size of teeth) were varied. For the cubes the starting layers were positioned at diﬀerent z positions to avoid simultaneous interaction of multiple parts with the recoater blade.Figure 5 shows the analysis results for the part area of a single part from job A.

The plot of A(x, z) shows larger elevations during the down-skin phase, 2.06 mm ≤ z ≤ 2.16 mm, and for increasing polar angle of the overhang (measured with respect to the vertical) in higher layers, z > 8.5 mm (Figure 5b). Most elevations are located at the part’s boundaries and not inside. Critical elevations appear only at the right edge and for large polar angle.These observations are supported by the elevation plots over z in Figure 5c, which show an initial bump at z = 2.0 mm (yellow circle) and a gradual increase for z > 8.2 mm. For z > 9.6 mm some layers are highlighted as Az(x) exceeds Acritical for some x positions (red triangles).To evaluate the impact of diﬀerent support structures in job B we analyze the part regions for three cubes in Figure 6. Support structures for parts STD35, STD36 and STD38 were(a)(b)(c)Figure 5: Analysis of overhanging geometry in x/z plot of elevated area. (a) part geometry with decreasing overhang angle for increasing z.

(a) part geometry with decreasing overhang angle for increasing z. (b) x/z plot of elevated area A(x, z), the starting z position and thereby the overhang angle of increasing elevations can be identiﬁed by analyzing the right part boundary. (c) the elevation statistics show the buildup of elevations for increasing z position. Note that the build job was aborted before the rectangular top part was built.0.120.121431570.00.28Area of Elevated Region [mm²]1591440.52.01410.00161X Position [mm]1.5151X Position [mm]0.28Area of Elevated Region [mm²]0.120.241.00.51492.00.002.5STD361482.5STD381471582.5STD351.00.201.51420.28Area of Elevated Region [mm²]1450.50.040.161500.00.000.080.041600.160.16146X Position [mm]1561.00.00.040.080.240.240.200.081.50.202.0Figure 6: Analysis of diﬀerent support structure conﬁgurations. Top: support conﬁgurations for connection to ﬁrst part layer.

Top: support conﬁgurations for connection to ﬁrst part layer. Bottom: elevations in the part region (as deﬁned by the part bounding box) for cubes from job B. Note that plots are cropped to z ≤ 2.5 mm to focus on the support structures.built with identical teeth height of 0.9, diﬀerent base length (0.6, 0.6 and 0.8) and varying top length (0.05, 0.25 and 0.15). The x/z plot shows almost no elevations in the support structures (z < 1.6 mm) for STD35, some elevations for STD36 and stronger elevations for STD38. Additionally, elevations in the ﬁrst part layers (z ≈ 1.7 mm) of STD35 and STD38 show a similar pattern, which is related to the regular structure of the supports. For STD36 In conclusion, the elevations are contiguous with two accumulation at the part corners.

For STD36 In conclusion, the elevations are contiguous with two accumulation at the part corners. the conﬁguration of STD35 achieves the highest process stability, while STD38 shows some critical elevations in the support structures, which may hinder usage in a batch job with multiple identical parts at the same x position.Two identical geometries from job B, which start at diﬀerent z positions are compared in Figure 7. Note that the plots of the part area start at the ﬁrst part layer, whose z position depends on the part placement inside the build job. The integrated x/z plots show more elevations in the down-skin layers (bottom) of the right part and an accumulation of elevated regions at the left part boundary for both parts.Additionally the relative elevated area Arel(x) was computed for each part and used to rank the parts according to their mean relative elevated area Arel,avg (Figure 8).

This enables quick assessment of a part’s process stability and can be used as a ﬁgure of merit during the job design phase.DiscussionThe color-coding of presented x/z plots is based on the critical area threshold Acritical = 0.1 mm2 deﬁned in [8], where Acritical was linked to an acceleration of the recoater blade of 9 m s−2, and uses a gradient color from white to orange for A(x, z) ≤ Acritical and from orange to red for Acritical < A(x, z) ≤ 3 · Acritical.

This coding was selected experimentally toFigure 7: Analysis of elevations inside the part for two identical geometries positioned at diﬀerent z positions.50002150010002500810002000150010Mean Relative Elevated Area [%]3500X position [px]3000130200052500STD30STD43STD36STD37STD28STD29STD44STD45STD35STD50STD51STD49STD34STD27STD46STD42STD48STD33STD38STD41STD47STD32STD39STD40STD3193000Y position [px]6475000STD30nameSTD36STD47STD42STD29STD329Mean Relative Elevated Area [%]STD44STD45STD41STD48STD39STD387STD27STD43STD28253STD49STD516STD35STD46STD4048STD50STD33STD371STD31STD34Figure 8: Relative elevated area plot. Analysis results for 25 identical cubes with diﬀerent support parameters from a single build job.

Analysis results for 25 identical cubes with diﬀerent support parameters from a single build job. Part regions are colored according to their mean relative elevated area Arel,avg, which enables quick assessment and comparison of process stability for single parts. Displayed labels are part names.visualize the diﬀerences in our build jobs and is not based on a comprehensive validation. Developing a validated color-coding could help to compare build job performance globally and avoid false alarms for stable jobs.While the ranking of parts based on mean relative elevated area establishes a deﬁnite order, the observed values are very low and the applied color scale could exaggerate the diﬀerences between parts (Figure 8).

An experimental validation and deﬁnition of boundaries for expected values under stable and unstable conditions will be performed in future work in order to provide recommendations for job stability optimization.ConclusionElevations of part regions pose a major risk to build reliability of laser beam melt- ing (LBM) processes as collisions between recoater blade and part may occur. Based on an image-based elevation detection method, we have presented quantitative measurement methods for analyzing part regions and parts for all layers of a build job. Visualization of elevated area A(x, z) over x (direction of the recoater blade) and z (layer position) enables quick comparison of stability for diﬀerent parts and localization of critical part regions, e.g. problematic geometrical features.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperElevated Temperature Mechanical and Microstructural Characterization of SLM SS304LG. R. Hecht*, S. P. Isanaka*, and J. W. Newkirk**Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409 **Department of Material Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractSLM built SS304L was annealed and water quenched to minimize residual stress andavoid carbide precipitation. Mini-tensile characterization of strength and elongation at temperature conditions up to 800˚C, along with observations of the associated microstructural transformations were utilized to understand the changes produced in SLM SS304L.

As-built and annealed specimens were found to exhibit decreasing strength and elongation with increasing temperature as expected. Carbide precipitates appeared after short times at high temperatures within both as-built and annealed specimens for all cases, but no brittle intermetallic phase development was observed for any of the temperatures investigated. While the lack of Sigma, Chi or Laves phases were anticipated, the premature formation of carbides is unexpected behavior for this composition of SS 304L. It is an indication of higher sensitivity of SLM made material. An additional change in the etch response was also observed between as-built and annealed specimens. It is theorized that annealing caused all ferritic and other residual phases present in as-built SLM SS 304L to fully transform into austenite.

It is theorized that annealing caused all ferritic and other residual phases present in as-built SLM SS 304L to fully transform into austenite. The cellular structure observed in the as-built specimen was also dissolved due to annealing and water quenching possibly leading to the strength loss observed.IntroductionAustenitic stainless steels provide excellent high-temperature corrosion resistance andmechanical strength of any alloy group, only bested by nickel-based alloys. For this reason, they are often the ideal choice for applications involving elevated temperatures and extreme oxidizing or reducing atmospheres where mechanical strength is a priority.

For this reason, they are often the ideal choice for applications involving elevated temperatures and extreme oxidizing or reducing atmospheres where mechanical strength is a priority. Austenitic stainless steels also exhibit good elongation at high temperatures, another very favorable characteristic, especially for applications where sudden failure is undesirable.At elevated temperatures, conventionally manufactured austenitic stainless steels areexpected to experience several phenomena leading to reduction in mechanical strength and ductility. As temperatures increase, the required energy for dislocations to slip past one another decreases, allowing plastic deformation to take place at increasingly lower values of applied stress leading to a reduction in mechanical strength. Even at low temperatures, thermal activation affects lattice resistance to dislocation glide.

Even at low temperatures, thermal activation affects lattice resistance to dislocation glide. Increased temperature results in a lower required temperature-assisted force to overcome obstacles to dislocation motion [1]. Chromium carbide may also precipitate at elevated temperatures if carbon within the material reaches levels ofThis work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.364supersaturation and diffusion rates sufficient for carbon and chromium to segregate into precipitates, causing a reduction in ductility and an increase in hardness [2]. If then subjected to an oxidizing atmosphere, these precipitates are then prone to rapid oxidation which can lead to the virtual disappearance of these metal carbides [3] also causing a reduction of mechanical strength and further reduction of ductility.

A third detrimental effect that may be seen in austenitic stainless steels at elevated temperatures is embrittlement due to the formation of a sigma phase. Sigma phase can form in austenitic stainless steels subjected to temperatures ranging from 500˚C to 980˚C after a few thousand hours in service under normal conditions, or almost instantly if the material has been cold worked [4].Due to limitations of conventional machining and manufacturing, selective laser melting (SLM) is an attractive alternative, allowing geometrically complex three-dimensional parts to be manufactured to a near end-use state, directly from computer-aided design files. Selective laser melting is a powder-bed fusion technique in which successive layers of pre-placed powder are bonded together with the use of a laser.

Selective laser melting is a powder-bed fusion technique in which successive layers of pre-placed powder are bonded together with the use of a laser. The low powder particle sizes and small input energy spot size in SLM allows parts to be made with high dimensional accuracy and thus require less post-processing when compared to other additive manufacturing techniques [5]. These advantages of increased design flexibility have been adopted by researchers across a range of conventional and novel materials to serve industries like aerospace and healthcare. For example, Concept Laser has produced bracket connectors to be used in the Airbus A350 XWB. This component was previously milled from aluminum but is now fabricated from titanium using an SLM process, reducing the component’s weight by over 30% [7].While SLM and other AM processes create three-dimensional geometries bymetallurgically bonding layers of metals and their alloys, they do so in a unique manner.

Owing to the small volume of phase change driven by high energy density, AM parts are composed of a complex framework of connected prior meltpools. While these meltpools can be observed microstructurally, their effect on AM-made materials at high temperatures warrant further investigation. Their cooling rates are also extremely high, leading to the possibility of creating unconventional microstructures. Before SS304L components manufactured using SLM can be used extensively at elevated temperatures, a thorough understanding of how the material behaves under these conditions is required. The goal of this research is to characterize the effects of elevated temperature on the mechanical and microstructural properties of SLM made SS304L. SLM made SS304L has been previously shown to exhibit high reactivity when aged at elevated temperatures when compared to its conventionally manufactured counterpart [6].

SLM made SS304L has been previously shown to exhibit high reactivity when aged at elevated temperatures when compared to its conventionally manufactured counterpart [6]. Thus, similar reactivity is expected to be observed at elevated temperatures.For the short-term tensile tests used for this characterization study, the main effect onmechanical properties at elevated temperatures is anticipated to be material softening. Due to the short time scale of the tensile tests, carbide precipitation and embrittlement due to sigma phase are very unlikely to be observed. For austenitic stainless steels with a carbon content less than 0.019% by weight, such as the samples used for this study as shown in Table 1, chromium- carbide is only expected to form after 100 hours at temperatures between 500˚C and 600˚C [2].

Sigma phase in austenitic stainless steels, again, only forms after a few thousand hours at temperatures ranging from 500˚C to 980˚C, with the exception of cold worked material for which formation begins instantly. Due to the complex stress-strain field present in SLM madeThis work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.365specimens, sigma phase could potentially begin forming in the investigated elevated temperature tensile tests. Simulations based on JMatpro indicate that sigma phase formation is unlikely for conventionally manufactured SS304L with the same chemical composition as the SS304L powder used within this study. Thus, sigma phase is not expected to form in either the as-built or annealed samples of SLM material exposed to high temperature as part of this research.

Thus, sigma phase is not expected to form in either the as-built or annealed samples of SLM material exposed to high temperature as part of this research. In order to confirm this, both as-built and annealed SLM SS304L samples have been characterized at temperatures ranging from 200˚C to 800˚C for changes in mechanical properties and microstructure. Samples were annealed at 1050˚C and water quenched to minimize possible residual stresses and to avoid carbide precipitation. In the past, researchers have reported high dislocation presence in SLM materials, but annealing and water quenching specimens is theorized to reduce dislocation density and consequently anticipated to increase ductility with the tradeoff of reduced strength in these SLM materials.Experimental ProcedureSpecimens used for this characterization study were built on the Renishaw AM250additive manufacturing system located at the Missouri University of Science and Technology.

The process parameters used for the fabrication of the specimens have been previously optimized to obtain maximum density and include a laser power of 200W, hatch spacing of 85μm, point distance of 60μm, and an exposure time of 75μm. It should be noted that the Renishaw AM 250 uses a point exposure scan strategy. The SS304L powder used for the fabrication of the specimens was purchased from LPW Technology. Table 1 displays the chemical composition of the powder.Table 1: Chemical Composition of SS304L Powder [9]Element Wt.C 0.018Element Al 0Wt.Mn 1.4 P 0.012Si 0.005 0.6 Pb 0SSn 0Cu 0.05 Co 0Ni Mo 0 9.8 B V 0 0Cb 0 O 0.02Cr 18.4 Ca 0Ti 0N 0.06Specimens were built as rectangular columns that were then used to make mini-tensile specimens from the interior of each block. This strategy was deployed to ensure that any effects of oxidation were not present on the tensile specimens and thereby avoided from the characterization study.

This strategy was deployed to ensure that any effects of oxidation were not present on the tensile specimens and thereby avoided from the characterization study. All cutting was performed using a Sodick Electrical Discharge Machine (EDM) to minimize any possible work hardening effects that other harsh machining processes would impart on the properties of the material. The use of mini-tensile specimens rather than an ASTM standard allows many data points to be acquired while minimizing the volume of material needed. This is ideal when dealing with SLM powder that is expensive due to the need for spherical particles with high purity [5] or in other situations where material costs and production volumes come at a premium.

Figure 1 displays a drawing of the mini-tensile specimens used for this study which have been developed at the Missouri University of Science and Technology and have a thickness of 1 mm [8].This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.366' I 4. 12 l_9.533 iJ 1 T7 IR0.87\ \ 6(106°'All d1mens1ons in mm1.19 -Figure 1: Dimensions of the Mini-Tensile Specimen Developed at the Missouri University of Science and Technology and Used for This Study. [8]Additionally, before the cutting of mini-tensile specimen, several rectangular columns were annealed by heating at 1050˚C for 30 mins before quenching in water.

[8]Additionally, before the cutting of mini-tensile specimen, several rectangular columns were annealed by heating at 1050˚C for 30 mins before quenching in water. This was done to avoid carbide precipitation and to minimize the residual stresses inherent to all SLM made specimens due to the rapid thermal cycling, phase change and directional building inherent to all metal AM processes. Prior to running tensile tests, all mini-tensile specimens where polished with 600 grit Silicon Carbide paper to remove any oxide formation from the EDM process.All of the tests were carried out on an Instron UTM using a custom built elevatedtemperature testing setup. An Applied Test Systems 3210 series tube furnace was fitted to the Instron UTM machine to allow specimens to be heated to temperatures of up to 1100˚C. Custom built pull rods were manufactured using the nickel based super alloy Inconel X due to its high temperature creep strength and low thermal conductivity.

Custom built pull rods were manufactured using the nickel based super alloy Inconel X due to its high temperature creep strength and low thermal conductivity. This ensured that high temperature tensile tests could be run for the full operational temperature range of the furnace while ensuring that the sensitive load cell remains well within the recommended calibrated temperature range for accurate measurements. Figure 2 displays the full elevated temperature set up used for this study.This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.367Figure 2: Elevated Temperature SetupA total of six short-term tensile tests each were run at room temperature, 200˚C, 400˚C,600˚C, and 800˚C for the as-built samples, and the annealed and water quenched samples.

Samples were held at temperature long enough to ensure uniform temperature and were then pulled at a constant strain rate of 0.06 mm/mm/s while the engineering stress-strain relationship was recorded. After completion of each test, the broken specimens were saved for further analysis.After completion of tensile tests, specimens were polished and etched with an electrolyticnitric acid setup (60% solution at 30V and 5mA for a 3 second exposure) to prepare for optical microscopy. Optical images were then taken of samples from each group of tests to analyze the microstructural characteristics of each.JMatPro was also used to simulate the effects of elevated temperature on the mechanical and microstructural properties of conventionally manufactured SS304L with the same chemical composition. A cell size of one micron was also used to simulate grain size and distribution in the software.

A cell size of one micron was also used to simulate grain size and distribution in the software. Intermetallic phase development and carbide precipitation were simulated for temperatures ranging from 100˚C to 800˚C at exposure times up to 1(cid:3400)1013 hours. The yield strength was simulated for temperatures ranging from 0˚C to 1400˚C. It should again be noted that JMatPro simulations are based on conventional manufacturing and therefore do not take into account the residual stresses, rapid cooling rates and other effects present in the SLM samples, and so some discrepancy is to be expected.This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.368Results and DiscussionThe results of the JMatPro simulation are displayed below in Figure 3.

Sigma and Lavesphase formation are predicted to form in this composition of SS304L, although neither are expected to within the temperatures or time frame of this characterization study. Sigma phase is predicted to appear after a few hundred to a few thousand hours at temperatures ranging from 400˚C to 750˚C while Laves phase formation takes nearly 500k hours beginning at around 300˚C. JMatPro predicts the formation of a M2 (C, N) precipitate but again only after a hundred hours of exposure to 750˚C. From these predictions, embrittlement due to the formation of these brittle intermetallic phases is unlikely within this characterization study of SLM SS304L.80 070 060 0u~ 50 0 ~ QJ ?400 QJ I-TTT Austenitic Stainless SteelSIGMA(0.1 %) ■ LAVE S(0 . 1 %) * M2(C ,N)(0.1 %) M(C,N)(0 1%)COMPOSITION (Wt%) Fe: 69.653 Cr: 18.4 Cu: 0.05 Mn: 1.4 Ni: 9.8 0 : 0,02 Si: 0.6 N: 0.

1 %) * M2(C ,N)(0.1 %) M(C,N)(0 1%)COMPOSITION (Wt%) Fe: 69.653 Cr: 18.4 Cu: 0.05 Mn: 1.4 Ni: 9.8 0 : 0,02 Si: 0.6 N: 0. 06 P: 0.012 S: 0.00 5 TRANS ITIONS (C) M23C6: not present CHI: not prese nt M(C ,N): 64 9.6 M2(C,N) 801 .5 LAVE S: 31 6. 1 M7C3: not present SIGMA: 659.5 M6C: not pres ent300200100 0100 0001 E7 Ti me (h )1 E91 E111E13Figure 3: Intermetallic Phase Development Simulated with JMatProJMatPro was also used to simulate the effect of elevated temperatures on the yield strength for comparison to the SLM SS304L material. As shown in Figure 4, the software predicts softening of the material with increasing temperature as expected. This behavior can also be confirmed with experimental data, although the physical tests indicate that the as-built specimens exhibit approximately 20 percent higher yield strength when compared to the simulated strength.

Simulations are based on conventionally manufactured materials so the discrepancy in the results could be a product of this assumption. Discrepancies could also be due to a higher dislocation density within the as-built specimens compared to the assumptions made within the simulation.

Before definitive conclusions can be made, further analysis using TEM is required to get an accurate idea of the dislocation density within the as-built specimen.This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.369High Temperature Strength500~ - - - - - - - - - - - - - - - - - -~400COMPOSITION (Wt%) Fe: 69.653 Cr: 18.4 Cu: 0.05 Mn: 1.4 Ni: 9.8 0 : 0,02 Si: 0.6 N: 0.06 P: 0.012 S: 0.005'i 300 ~ "' "' .,b 200 (fJ■ 0.2% Proof Stress100200400600800100012001400Temperature(C)Heat treatment : 820.0 C AUSTENITE 99.98% Matrix grain size (microns) Fixed strain rate: 2.SE-4 (1/s)Figure 4: Yield Strength Simulated with JMatProFigures 5, 6, 7, and 8 show the experimental results acquired for the as-built and the annealed specimen.

As can be seen, the annealed specimens exhibit lower strengths compared to the simulated and as-built data. Inversely, the annealed specimens exhibit much higher elongation at elevated temperatures with remarkable stability for the full range of temperatures investigated. This behavior could be attributed to the dislocation density possibly being reduced during annealing. Additional microstructural transformations could also be contributing to this mechanical performance behavior.SLM made specimens have been known to possess higher yield strengths when comparedto their wrought and cast counterparts while wrought and cast characterization studies are typically used as the basis for the JMatPro simulations. This is confirmed in this study, with the experimental yield strengths of SLM specimens being higher than the predictions from JMatPro.

This is confirmed in this study, with the experimental yield strengths of SLM specimens being higher than the predictions from JMatPro. Interestingly, the trends from both the as-built and the annealed strength investigations in Figure 5 match the prediction of conventionally manufactured SS 304L from JMatPro. This implies that the SLM material behaves similar to its conventional counterparts when exposed to high temperatures.

This implies that the SLM material behaves similar to its conventional counterparts when exposed to high temperatures. Were it not for the starting yield strength being higher in the SLM specimens, the trend might match the JMatPro predictions.This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.370600500..400;.~ 300 "' >200100--------rTI 600500..400m~ 300 "' >200100.......02120040060080021200400600Temperature (c)Temperature (c)Figure 5: As-built (left) and Water Annealed (right) Yield Strength.700 --800.. ~ 500 ... "' :,600;.400---r800700 -.. ~ 500 ...

As can be seen, the etchresponses of the as-built and annealed specimens were dramatically different.Room Temp.200˚C400˚C600˚C800˚CFigure 9: As-Built Specimen Etched with Electrolytic Nitric Acid Setup Under 5x Magnification.Room Temp.200˚C400˚C600˚C800˚C200 µmFigure 10: Annealed Specimen Etched with Electrolytic Nitric Acid Setup Under 5x Magnification.This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.372Within the 100x magnification images of the as-built specimens shown in Figure 11, intermetallic phase development seems to be absent from both the specimens tested at room temperature and 800˚C. Unexpectedly, carbide precipitates do appear to have formed within the as-built specimen tested at 800˚C.

Unexpectedly, carbide precipitates do appear to have formed within the as-built specimen tested at 800˚C. The absence of other intermetallic phases agrees with literature and the JMatPro simulations for the short time scale of short-time tensile tests, but the formation of the carbides does not. The loss of strength seen in the experimental characterization could be attributed to the dissolution of the material’s cellular structure when exposed to high temperatures, but this also does not seem to be the case of the as-built specimens.Room Temp.800˚CCarbidesFigure 11: As-Built Specimen Etched with Electrolytic Nitric Acid Setup Under 100x Magnification.Carbide precipitation was also observed within the annealed and water quenchedspecimen tested at room temperature as well as at 800˚C as shown in Figure 12. Intermetallic phase development seems to absent from the annealed specimen for both cases shown below.

Intermetallic phase development seems to absent from the annealed specimen for both cases shown below. Carbide precipitation was not predicted to occur for either as-built or annealed specimens and is theorized to have been accelerated by the SLM process. Other researchers have found that SLM materials are more sensitive than their conventionally manufactured counterparts and this premature precipitation of carbides could be further indication of that behavior.

In both these cases, dislocation slip at high temperatures could be one of the reasons that contribute to the mechanical performance of these materials.Room Temp.800˚CCarbidesCarbidesI II10µmFigure 12: Annealed Specimen Etched with Electrolytic Nitric Acid Setup Under 100x Magnification.This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.373As stated previously, the etch response of the annealed specimens was dramaticallydifferent compared to that of the as-built specimens. Although the electrolytic nitric acid etch procedure was exactly the same, the phase resolution in the annealed case is not apparent. Due to the specimens being SLM made, it is expected to see track and meltpool boundaries in the microstructure.

Due to the specimens being SLM made, it is expected to see track and meltpool boundaries in the microstructure. The cellular structure which is present in the as-built SLM specimen is also not apparent in the annealed and water quenched material. It is theorized that annealing the specimens has resulted in a transformation of any ferrite or other residual phases present into austenite, as this could explain the drastic change in etch response compared to as-built specimen. Also, the dissolution of the cellular structure could explain some of the differences within the microstructural images. This could also explain the change in mechanical performance between the as-built and annealed specimen, as the dissolution of the cellular structure could produce a slight loss in strength and a corresponding increase in ductility.ConclusionThe mechanical and microstructural properties of as-built and annealed SLM SS304Lhave been investigated at elevated temperatures up to 800˚C.

Both as-built and annealed specimens exhibited a decline in strength and ductility as temperatures increased. Intermetallic phase development appeared to be absent within both as-built and annealed specimens for all temperatures investigated. Interestingly, carbide precipitates did appear to be present within the as-built specimens tested at 800˚C, as well as in all of the annealed specimens. Thus, the decline in strength and elongation as temperature increases is theorized to be due to dislocation slip as elevated temperatures reduce the required energy for dislocations to slip past one another, coupled with the loss of the cellular structure in the annealed specimen. Carbide precipitation could also play a role in the decrease in ductility at elevated temperatures. Annealed specimens were observed to exhibit a much different etch response as compared to the as-built specimen.

Annealed specimens were observed to exhibit a much different etch response as compared to the as-built specimen. It is theorized that the process of annealing the specimens has resulted in ferritic and other residual phases present to transform fully into austenite as well as the dissolution of the cellular structure.Future WorkFurther investigation is planned to better understand what has been observed during thischaracterization study. Dislocation studies will be performed employing transmission electron microscopy (TEM). Electron backscatter diffraction (EBSD) will be used to gain a thorough understanding of the microstructural phase transformations observed and the grain size present within the specimens.

Electron backscatter diffraction (EBSD) will be used to gain a thorough understanding of the microstructural phase transformations observed and the grain size present within the specimens. Finally, further optical microscopy is underway to investigate grain boundary migration and realignment that may have taken place.Special ThanksA very special thanks goes out to Alexander Lipari for all of our microscopy imaging, Adam Fleischman for help completing our elevated temperature tensile tests, and Kyle Stagner for preparing all of our samples.This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.374References[1]Courtney TH (2013) Mechanical Behavior of Materials, pp. 298, McGraw Hill Education (India), New Delhi[2]Kelly, J. (2006). “Stainless Steels.” Mechanical Engineers Handbook, pp.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperElevated temperature microstructure stability of SLM 304L stainless steelTarak Amine and Joseph Newkirk Materials Science and Engineering Department, Missouri University of Science and TechnologyAbstractAt elevated temperatures changes in metallurgical structure can be expected for almost any steel or alloy. In stainless steels, the changes can be grain growth, carbide precipitation, ferrite decomposition, or embrittlement. These phenomena can significantly affect the properties of the stainless steel and would potentially change the functionality of the component. Therefore, the impact of elevated temperatures on the microstructure of SLM 304L stainless steel was studied.

Therefore, the impact of elevated temperatures on the microstructure of SLM 304L stainless steel was studied. The work reported here investigates the influence of different aging times at 300C on microstructure and mechanical properties of stainless steels (304L) fabricated with the Selective Laser Melting (SLM) process. Microstructure and mechanical properties were dramatically effected at temperatures much lower than expected when compared to samples of wrought stainless steel. The stainless steel fabricated using the SLM process was more kinetically active than expected based on previous studies of 304L. Results of this study will be presented along with possible reasons for the higher activity.1. IntroductionAdditive manufacturing materials fabricated by the SLM process are used increasingly more in industrial applications due to SLM’s ability to fabricate complicated parts geometry coupled with restricted tolerances and the availability of a wide range of alloy compositions.

Moreover, austenitic stainless steels are extensively used in a lot of applications where operating temperatures are in the range 290–(cid:22)(cid:21)(cid:24)(cid:3)(cid:405)(cid:38), such as the nuclear industry for core internal structures and primary recirculation piping. This is due to its superior properties in corrosion resistance, ductility, strength, formability, and weldability [1].

However, the ferrite phase is a metastable phase, which transforms to various other embrittling phases at elevated temperature applications [2–5]. Therefore, the ferrite phase may become embrittled after exposure to long term aging at low aging temperatures.The two ferrite and austenite constituents cannot be equally stable at elevated temperatures. It is anticipated, therefore, that the austenitic stainless steel of the alloys can create the different potential for thermal aging embrittlement in different portions through decomposition of existing phases and/or precipitation of new phases. In this work, low temperature aging behavior of austenitic SS type 304L was studied after aging at 300 (cid:405)C for different aging time up to 1000 hr.2.

In this work, low temperature aging behavior of austenitic SS type 304L was studied after aging at 300 (cid:405)C for different aging time up to 1000 hr.2. Materials and Test Method©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.640In this work, 304L stainless steel powder with particle sizes less than 45 (cid:151)m were used as the starting material in SLM experiments. The chemical composition of the powder is listed in table 1. A Renishaw AM 250 machine was used to process the 304 stainless steel powder. One sample was investigated in the as-built condition. The other samples were aged at 300 ºC for different aging time after fabrication (20, 45 minutes, 25, 250, 500, and 1000 hr). The aged samples were packed in heat treating bags, made of type 309 stainless steel, to reduce decarburization and oxidation.

The aged samples were packed in heat treating bags, made of type 309 stainless steel, to reduce decarburization and oxidation. Then the samples were cooled in air after the aging process. Once cool, samples were unpacked and prepared for tensile testing.Table 1:- The chemical composition of 304L stainless steel powder Element Cr Wt%Cu Mn Ni 9.9 1.4 0.1O 0.02Si 0.63C 0.01518.5N 0.09P 0.012S 0.004 BalanceFeThe samples were fabricated to miniature test pieces in the shape of a “dogbone” and was designed developed by ABI services LLC. The total sample height was 16 mm. The samples were subsequently sliced to 1.5 mm thickness by using Electrical Discharge Machining (EDM). The kerf loss attributed to the EDM cutting was 0.5 mm. Six (6) samples could be extracted from the total 12 mm height. The samples were numbered 1 through 6, starting with the top surface as shown in Figure 1.

Six (6) samples could be extracted from the total 12 mm height. The samples were numbered 1 through 6, starting with the top surface as shown in Figure 1. All samples were polished sequentially with 400, 600 and 800 grit sandpapers prior to tensile testing.1 2 3 4 5 6Figure 1: Sliced mini tensile samples of 1.5 mm thickness. Numbers indicate the sample location in term of build height (1 is top of the build, 6 is bottom of the build).The Automated Ball Indentation (ABI) technique provides a nondestructive and localized alternative to the destructive tensile test [6]. The ABI indentation technique requires a planar spacing of the indents that must be a minimum of three diameters from their centers and be at least two diameters from free edges.

The ABI indentation technique requires a planar spacing of the indents that must be a minimum of three diameters from their centers and be at least two diameters from free edges. The indenter diameter used was 0.03 inch and the load cell used©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.641had a 250 lb capacity. Five ABI measurements were taken on the fabricated blocks of 304L for each of the as-built and aged conditions. Figure 2 shows a micrograph of the general appearance of the indent made in as-built and aged samples using a 0.03 inch diameter indenter from an eight- cycle automated indentation.Figure 2: ABI indent in AM304L using a 0.03 inch diameter indenter in as-built material. Note indent covers multiple laser tracks.Electrolytic etching was used for etching the specimens.

Note indent covers multiple laser tracks.Electrolytic etching was used for etching the specimens. The microstructure of the fabricated parts was investigated by optical microscopy. The phase analysis was conducted by X-ray diffraction analysis (XRD) using Philips X Pert-MPD System with Cu Ka radiation and 40 kV and 30 mA operating conditions. Also, initial thermal analysis study has been done by using Differential Scanning Calorimetry (DSC). A microhardness tester equipped with Vickers pyramid indenter was used for microhardness measurements, which were randomly measured on the fabricated samples. A 100 g load was applied by the indenter for 10 seconds.3. Results and Discussion3.1.Microstructure of AM 304LFigure 3 shows optical micrographs at different locations of the fabricated sample part of as-built material.

Results and Discussion3.1.Microstructure of AM 304LFigure 3 shows optical micrographs at different locations of the fabricated sample part of as-built material. The sample was etched by electrolytic etching with a solution of 60 percent nitric acid and 40 percent water, and with a relatively low voltage of 1 volt, to delineate the grain boundaries. The optical micrographs indicate that complete melting of the powder occurred. Irregularly shaped pores appeared at various locations. The etching procedure delineated the grain boundaries. The laser tracks, hatching between two neighbor layers, and the different scan strategy©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.642at the fabricated build edges clearly can be seen.

The microstructure of as-built 304L material consisted of a mixture of austenite and (cid:303)-ferrite as verified by X-ray diffraction.Build edgeDeformed areaFigure 3: Optical micrographs at different location of As-built material. Locations as marked on broken tensile samplesFigure 4 shows optical micrographs at different locations of the aged samples. The aging temperature was 300 ºC and the aging time was 250hr and 1000hr. Several pores appeared at the hatch boundaries. The optical micrographs reveal a distinct region in the microstructure on 300 °C aged samples for 250, 500, and 1000 hr at the grain boundaries.

Several pores appeared at the hatch boundaries. The optical micrographs reveal a distinct region in the microstructure on 300 °C aged samples for 250, 500, and 1000 hr at the grain boundaries. This type of structure could be carbide formation (M 23C6 or M 2C).©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.643(a)Pore between boundaries(b)Figure 4: Optical micrographs at different location of Aged samples at 300 ºC for (a) 250hr, (b)1000hrFigure 5a shows a typical backscatter SEM image of the cellular structure present in the sample.

The backscatter mode shown in the higher magnification image in figure 5b reveals two distinct©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.644regions in the microstructure denoted by 1 and 2.

This has not been verified here.During long-term exposure to elevated temperatures of austenitic stainless steels some transformations involving the precipitation of sigma, chi, intermetallic phases, ferrite, and various carbides in addition to change in ferrite / austenite content and alloy recrystallization can take place. These phases are often found at grain boundaries and other high-energy intragranular sites in the form of cubic or needle-like precipitates [12, 13]. These transformations induce remarkable changes in mechanical properties of aged austenitic stainless steels.According to JMatPro software calculation based on cast parts, these phases cannot be formed at low temperatures, such as 300ºC until years of temperature exposure as shown in figure 6. The JMatPro software assumes a conventional casting cooling rate for the creation of the microstructure, which should be significantly slower than for AM.

The JMatPro software assumes a conventional casting cooling rate for the creation of the microstructure, which should be significantly slower than for AM. The JMatPro software calculations were based on 0.1% of amount transformed.

DSC measures the energy absorbed or released by the sample when it is heated or cooled. The sample was linearly heated twice to 600 ºC at a heating rate of 10 ºC/min, as shown in figure 8. The DSC curves were not identical for the two runs, demonstrating that the first heated run has a significant effect on the additive material thermal behavior. This could indirectly indicate the magnitude of the energy stored in the as-built material.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.648100First runSecond runFigure 8:- DSC curve for additive manufactured SS 304L3.2.Mechanical PropertiesTensile tests of SLM stainless steel 304L was carried out in the as-built and aged conditions. Miniature flat tensile specimens having a cross-section of 1.0 mm × 1.5 mm and a gage length of 8.0 mm.

Miniature flat tensile specimens having a cross-section of 1.0 mm × 1.5 mm and a gage length of 8.0 mm. The tensile samples were preloaded with 50 lb and the crosshead velocity was set at 0.0005 in/sec. The load cell used had a 1000 lb capacity.Figure 9 shows the tensile test results of the as-built material. The results show that the sample from the top of the build is stronger but less ductile compared to the other samples. The other samples have relatively the same strength but slightly different ductility (however, there is no apparent trend in ductility). The sample at the top of the build (sample #1) has the largest strength but the smallest ductility. The top of the build was the last to be melted and did not go through a reheat cycle. Hence it was subjected to the least heat and expected to be harder and less ductile than the other samples from the middle and bottom of the build.

Hence it was subjected to the least heat and expected to be harder and less ductile than the other samples from the middle and bottom of the build. The other samples were subjected to more heat, having been reheated because of their location in the build.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.649Figure 9: Engineering stress–strain curves of as-built material samples at different reference positionsFigure 10 shows the comparison of the tensile strength of the as-built and aged samples with different aging parameters.

The results show that both the yield strength and ultimate tensile strength (UTS) decreased after aging.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.650YSUTS800750700650) a P M( h t g n e r t S600550500450400350300250200150100As-BuiltAged 300°C, 25hr Aged 300°C,250hr Aged 300°C,500hr Aged 300°C,1000hrMaterial ConditionFigure 10: Comparison of tensile strength of the as-built and aged samples with different aging parametersFigure 11 shows a comparison of tensile strength results from the ABI indentation measurements of the as-built and aged samples at different reference positions. The results show that ultimate tensile strength (UTS) decrease in the aged samples. The ABI results show higher strength values compared to the mini-tensile results.

The results show that ultimate tensile strength (UTS) decrease in the aged samples. The ABI results show higher strength values compared to the mini-tensile results. The reason for the difference could be porosity, which the ABI indentation technique is less sensitive to since the ABI test is compressive and multi-axial in nature while the tensile test interrogates a much larger test volume and is an axial (single-axis) test.

However, the ABI still shows relatively good agreement with the mini tensile results of yield and ultimate tensile strengths.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.651YSUTS800700600) a P M( h t g n e r t S500400300200100As-BuiltAged 300°C, 25hr Aged 300°C, 250hr Aged 300°C, 500hr Aged300°C, 1000hrMaterial ConditionFigure 11: Comparison of tensile strength from ABI measurements of the as-built and aged samples with different aging parametersThe microhardness test results on the as-built and aged samples at 300 ºC for 20, 45 minutes, and 250, 500, and 1000 hours are shown in figure 12. The load was 981.2 mN and pressing time was 10 secs. The microhardness values shown are the average of ten measurements.

The load was 981.2 mN and pressing time was 10 secs. The microhardness values shown are the average of ten measurements. Despite some variation in the measured values, a clear trend of continuously decreasing average of the microhardness with increasing aging time was observed.

The other samples have relatively the same strength and slightly different ductility (there is no apparent trend in ductility).2- The aged samples have relatively the same strength at each location and slightly differentductility (there is no apparent trend in ductility)3- Tensile results show that both UTS and YS decreased with increasing aging time. 4- The ABI shows good agreement with the mini tensile results for both the yield and ultimatetensile strengths.5- The microhardness decreases with increased aging time. This could be due to the decrease(cid:76)(cid:81)(cid:3)(cid:87)(cid:75)(cid:72)(cid:3)(cid:303)-ferrite content and energy released.6- The (cid:68)(cid:88)(cid:86)(cid:87)(cid:72)(cid:81)(cid:76)(cid:87)(cid:72)(cid:3)(cid:68)(cid:81)(cid:71)(cid:3)(cid:303)-ferrite phases were observed in both as-built and aged samples.

7- The as-built material appears to have a higher energy stored in the microstructure and ismore kinetically active than wrought material of the same composition.©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.653AcknowledgmentsThis work has been funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.References[1] A.J. Sedricks, Corrosion of Stainless steels, 2nd ed., John Wiley and Sons, New York, NY, 1996, p. 13. [2] J.M. Vitek, S.A.

[12] Spruiell, J. E., Scott, J. A., Ary, C. S., & Hardin, R. L., “Microstructural stability of thermal- mechanically pretreated type 316 austenitic stainless steel,” Metallurgical Transactions, 4(6), 1533-1544 (1973). ©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.654[13] K.H. Lo, C.H. Shek, J.K.L. Lai, “Recent developments in stainless steels,” Mater Sci Engineer R (65) 39-104 (2009).©2016 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.655.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEmbedding Fiber Optic Sensors in Metal Components via Direct Energy DepositionElias B. Snider*, Rony Kumer Saha+, Cesar Dominguez †, and Jie Huang+, Douglas A. Bristow*Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65401 +Department of Electrical Engineering, Missouri University of Science and Technology, Rolla, MO 65401 †Los Alamos National Laboratory, NM 87545AbstractFiber optics are useful as strain and temperature sensors in a variety of applications involving high-value parts. Embedding fiber optic sensors into end-use parts can allow for real- time strain and temperature monitoring of these parts in extreme conditions.

Embedding fiber optic sensors into end-use parts can allow for real- time strain and temperature monitoring of these parts in extreme conditions. Direct energy deposition processes have distinct advantages for producing parts in high-value embedded sensing applications, however, ensuring survival of the optical fiber during embedment is challenging. A method of fiber embedment using high-temperature ceramic adhesives is presented as a flexible method of embedding fiber optic sensors into end-use parts such as to preserve fiber transmission and sensor integrity. Example specimens are presented, and functionality of sensing capabilities is demonstrated.I.IntroductionOptical fibers are useful as strain and temperature sensors and have gained significantattention in structural heath monitoring especially in large civil and infrastructural applications [1, 2, 3].

Recent interest has developed in using optical fibers in embedded sensing applications for smaller-scale metal parts to monitor strain and/or temperature distributions during use in engineering applications. Additive manufacturing processes are well-suited to embedding optical fibers in that they can deposit material around or over optical fibers. As a result, optical fiber sensors can be placed inside a part, allowing for more detailed strain and temperature information.

As a result, optical fiber sensors can be placed inside a part, allowing for more detailed strain and temperature information. In addition, by using optical frequency domain reflectometry (OFDR), a sensing technique capable of determining distributed strain measurements along the length of an optical fiber, strain distributions and concentrations can be determined continuously over a fiber sensor embedded in a part.Three main additive manufacturing methods, each with distinct advantages and disadvantages in part applications, have shown promise for integrating fiber optic sensors into metal parts: ultrasonic additive manufacturing (UAM), selective laser melting (SLM), and directed energy deposition (DED). Significant demonstrations of embedded optical fibers have been performed on ultrasonic additive manufacturing (UAM) machines specifically with aluminum [4, 5, 6] and stainless steel [7].

Significant demonstrations of embedded optical fibers have been performed on ultrasonic additive manufacturing (UAM) machines specifically with aluminum [4, 5, 6] and stainless steel [7]. Embedded fiber sensors with electroplated metal coatings have been incorporated into SLM parts made with stainless steel and other metal alloys [8, 9, 10]. DED embedment attempts have also produced embedded fiber sensors in titanium and Inconel alloys [11, 12] by employing substantial metal coatings on the optical fibers.1070This work explores the use of a high-temperature adhesive to embed optical fibers withinDED additively manufactured stainless steel parts and demonstrates the practical advantages of using this method as opposed to the use of significant metal coatings in embedded fiber sensors.II.

Metal-Coated Embedded Fiber SensorsMany of the efforts toward embedded fiber sensors in DED and SLM additive manufacturing have used electroplating methods to deposit very large nickel or copper coatings onto the fiber sensors [8, 11, 12]. These coatings serve to protect the fiber from laser exposure during the embedment process and allow for direct metal-to-metal fusion of the fiber coating to the deposition metal. The process for embedding nickel-coated fiber (Figure 1) begins by electroplating either copper-clad or bare silica fiber [13] to relatively large diameters (350μm – 2mm compared to the fiber core diameter of 125μm) as shown schematically in Figure 1.a. A base material is printed using the additive manufacturing process (Figure 1.b) and a slot is either incorporated into the geometry of the print or machined into the material (Figure 1.c). The metal- coated fiber sensor is then placed into the slot (Figure 1.d).

The metal- coated fiber sensor is then placed into the slot (Figure 1.d). Lastly, printing resumes, thereby consolidating the fiber into the material of the part (Figure 1.e).Figure 1: Fiber Sensor Embedment Procedure with Nickel CoatingsThe electroplating process can be time-consuming and difficult, often resulting insignificant variations in the diameter and straightness of the metal coatings. Because of these variations, the slot often requires oversizing to accommodate the fiber (Figure 2). Further, the final coated metal fiber is easily susceptible to damage if significantly bent or deformed.

Further, the final coated metal fiber is easily susceptible to damage if significantly bent or deformed. As a result, the optical fiber must be electroplated into the geometry of the final application, which presents significant challenges if the application geometry is not relatively straight.1071Figure 2: Example of Oversizing Fiber SlotIt has also been observed that embedding nickel coated fibers with DED or SLM additive methods results in limited bonding to the fiber; because of the constraints of these processes, the deposition material only bonds to the top surface of the fiber, leaving a cavity beneath the fiber that is not bonded to any of the surrounding material. This phenomenon is worsened with processes that have coarser resolution as it forces a deeper slot with less precise geometry unless the slot is machined.III.

This phenomenon is worsened with processes that have coarser resolution as it forces a deeper slot with less precise geometry unless the slot is machined.III. Adhesive-Embedded Fiber SensorsNoting the difficulties involved with securing and embedding nickel-plated fiber sensors, an alternative method was developed using high-temperature adhesives to allow more flexibility in embedment design and reduce some of the constraints imposed by the larger nickel coatings. The cross-sections of the embedment process are shown in Figure 3. This process begins with securing the same copper-clad optical fibers (Figure 3.a) with the exception that the fiber is not electroplated with nickel. The base material is printed (Figure 3.b) in the same way as in Figure 1.b. A slot is again either incorporated into the path planning of the build or machined into the existing material (Figure 3.c); however, the dimensions of the slot may now be considerably smaller.

A slot is again either incorporated into the path planning of the build or machined into the existing material (Figure 3.c); however, the dimensions of the slot may now be considerably smaller. A high-temperature adhesive is then injected into the slot (Figure 3.d) and the fiber sensor is immediately placed into the slot while the adhesive sets (Figure 3.e). Before the adhesive has set, a sacrificial plug is placed over the open interface of the slot (Figure 3.f). This sacrificial plug should be of a similar or compatible material to the deposition material such that it will fuse well to the deposition material and serves to both protect the fiber and adhesive as well as to act as a bonding medium between the deposition and the adhesive. Additionally, incorporating a bevel into the slot aids in placement and seating of the plug onto the surface of the slot.

Additionally, incorporating a bevel into the slot aids in placement and seating of the plug onto the surface of the slot. Material deposition then resumes, consolidating these components into the part (Figure 3.g) and yielding a finished part capable of embedded sensing.1072Figure 3: Fiber Sensor Embedment Procedure with AdhesiveThis method of using an adhesive to embed the fiber sensors overcomes several of thechallenges inherent to using large nickel coatings for embedded fiber sensors. Firstly, it eliminates the need to electroplate large amounts of nickel onto the fibers. Small-diameter copper coated fibers are commercially available and have less variance in diameters as compared to the large-diameter nickel coated fibers.

Small-diameter copper coated fibers are commercially available and have less variance in diameters as compared to the large-diameter nickel coated fibers. In addition, small-diameter copper coated fiber is very flexible and much more robust to bending than nickel coated fibers, allowing for the design of more complex embedment geometry.In addition, the slot may be significantly smaller than with the previous method resulting from the reduction in diameter and removal of the oversizing constraints. This allows for a much smaller footprint in the part and more precise placement of the fiber, especially if the slot is machined rather than incorporated into the build path. Lastly, the fiber is adhered on all sides by the adhesive, eliminating the void below the fiber sensor and enabling the fiber to be fully bonded. The adhesive is bonded to all three surfaces of the slot and is also bonded to the sacrificial plug which is in turn fused to the deposition material.IV.

The adhesive is bonded to all three surfaces of the slot and is also bonded to the sacrificial plug which is in turn fused to the deposition material.IV. Experimental SetupUsing the proposed method, an ASTM E8/E8M subsize tensile specimen was produced using Laser Engineered Net Shaping (LENS©) Print Engine and 53-150µm stainless steel 316L powder. The initial deposition was oversized such that machining would bring the specimen to exact ASTM geometry. The adhesive embedment procedure was followed with the slot (Figure 3.c) being incorporated into the build path planning (as opposed to being machined after initial deposition). After embedment (Figure 3.g), the specimen was separated from the build plate with1073a wire EDM and machined to precise geometry according to the ASTM specifications with rectangular cross-section and a thickness of 5.74mm (Figure 4).

A single mode, copper-coated optical fiber was used with a core diameter of 9μm, a cladding diameter of 125μm, and a copper coating thickness of 15μm. The adhesive used was a zirconia-based ceramic adhesive and is rated up to 2200 °C, ResbondTM 904. Lastly, a 316L stainless steel drawn wire was used as the sacrificial plug material.Figure 4: Embedded Fiber Tensile SpecimenThe strain sensing capabilities of this method were verified by performing an axial straintest in which the tensile specimen was loaded into a Tinius Olson T25 running Horizon control software with a 5kN load cell and subjected to a set of small deformations over approximately 15-30% of the elastic region of the tensile specimen (though this is difficult to determine as the specimen did not undergo any thermal treatment and as-deposited material characteristics can be unpredictable). The termination end of the sensor was adhered to the machine to prevent motion and subsequent noise in fiber strain measurements.

The termination end of the sensor was adhered to the machine to prevent motion and subsequent noise in fiber strain measurements. The strain was monitored during the tensile test using an Epsilon® Model 3542 Axial Extensometer and recorded in time. The distributed strain data from the fiber sensor over the 20mm extensometer section was similarly recorded using a Luna ODiSI 6000 OFDR system sampled at 1 Hz with a spatial resolution of 0.65mm and the resulting elongation over the extensometer section was calculated. This loading was applied in two separate tests for the specimen. Figure 5 contains the tensile test experimental setup.1074Figure 5: Tensile Test SetupV. ResultsA sample of the distributed strain profile of the optical fiber in the interest region (tensilespecimen) at the final loading point of Trial 1 is shown in Figure 6. Note that the ideal strain profile would be uniform in the gauge length and that the extensometer region is smaller than the full gauge length.

Note that the ideal strain profile would be uniform in the gauge length and that the extensometer region is smaller than the full gauge length. In calculating elongation (Figure 7, Figure 8, and Figure 9), only the fiber data in the test region was considered as it is the only direct comparison to the extensometer data (expressed as a percent change in distance between extensometer blades). Also note that the strain profile is not to scale; it is included only for a generalization of the shape of the expected strain profile. As discussed in Figure 10, the ideal strain magnitude would be greater than is shown in Figure 6. Some noise is present in the gauge length of the test specimen and is likely due to misalignments of the optical fiber within the slot. Because this slot was not machined to precise tolerances, the slot was not tight against the fiber, allowing for some amount of variation of the fiber placement within the slot.

Because this slot was not machined to precise tolerances, the slot was not tight against the fiber, allowing for some amount of variation of the fiber placement within the slot. If the fiber was not exactly straight when the adhesive set, any curvature in the fiber results in a misalignment from the axial strain applied to the fiber and would attenuate the strain signal slightly. This is likely the cause of the variations observed in the gauge length and precision-machining a slot with tighter tolerances would allow for more precise control over the placement of the fiber sensor if this were to be of concern in future embedment procedures.1075Figure 6: Sample of Strain Profile from End of Loading in Trial 1Results of each trial are shown in Figure 7 and Figure 8, and are shown together in Figure 9.

While the strain transmission to the fiber is dampened, the correlation is relatively linear and is smooth.1076Figure 7: Trial 1 of Axial Strain with Adhesive-Embedded Fiber SpecimenFigure 8: Trial 2 of Axial Strain with Adhesive-Embedded Fiber SpecimenIt is interesting to note that the extensometer measurement varies more over the twoloading cases in Figure 9 (which should have been identical barring any machine disturbances) and is likely due to misplacement of the extensometer or result of some other experimental disturbance. The fiber strain loading from each trial appears to be very repeatable and is nearly identical for the two trials.Figure 9: Axial Strain Trials with Adhesive-Embedded Fiber Specimen1077Figure 10 shows the sensitivity of the fiber sensor over the two loading trials. Thevariation in the first few sample points is almost certainly due to numerical error and sensor resolution limitations as the elongation approaches zero.

Thevariation in the first few sample points is almost certainly due to numerical error and sensor resolution limitations as the elongation approaches zero. The fiber sensor strain measurements are attenuated to approximately 20-30% of the true strain values and is likely due to the translation of strain through the adhesive, copper cladding, and fiber cladding (a silica coating over the fiber core). While the strain is attenuated, the sensor is still quite capable of measuring even very minute strains (close to 1μm) with calibration.Figure 10: Sensitivity of Fiber Sensor to Actual StrainVI. Conclusion and Future WorkOptical fibers are useful for a variety of sensing applications and open opportunities for revolutionary embedded strain and thermal sensing capabilities in end-use parts.

Conclusion and Future WorkOptical fibers are useful for a variety of sensing applications and open opportunities for revolutionary embedded strain and thermal sensing capabilities in end-use parts. The flexibility of the process presented in this work retains the flexibility of distributed strain sensing by eliminating the need for substantial metal coatings while also reducing the difficulty in obtaining the materials necessary to produce a part with embedded fiber sensors. In addition, the geometrical impact of the sensor on the part may be lessened, especially with precision- machined slots, and the void and unbonded regions are mitigated. While strain readings are attenuated, these embedded sensors are still capable of measuring very minute strains.

While strain readings are attenuated, these embedded sensors are still capable of measuring very minute strains. Further, by using precision machined slots, the reduction of strain transmission through the adhesive may be lessened to promote a more sensitive strain reading from these embedded sensors.The development of optical fiber embedment methods is still in its infancy and end-usecases for metal parts with embedded sensors are largely unexplored. Future work should develop the methods necessary to reliably embed and more rigorously characterize the strain readings from these sensors with as little impact on the part as possible. Machining fiber slots should be explored for its advantages in accuracy of fiber placement, especially as the toolpaths for these slot geometries could be more accurately correlated to CAD geometry for viewing strain and temperature data on a virtual representation of the part.

Various adhesives may also grant better material properties and strain transmission to the fiber. End-use parts with complex geometries would be well-suited for embedded fiber sensing applications and should also be explored in the interest of furthering this work into applications.1078VII. References[1] H.-N. Li, D.-S. Li and G.-B. Song, "Recent applications of fiber optic sensors to healthmonitoring in civil engineering," Engineering Structures, vol. 26, no. 11, pp. 1647-1657, 2004.[2] J. Ko and Y. Ni, "Technology developments in structural health monitoring of large-scalebridges," Engineering Structures, vol. 27, no. 12, pp. 1715-1725, 2005.[3] L. S. M. Alwis, K. Bremer and B. Roth, " Fiber Optic Sensors Embedded in Textile-Reinforced Concrete for Smart Structural Health Monitoring: A Review," Sensors, vol. 21, no. 15, p. 4948, 2021.[4] X. L. He, Z. Q. Wang, D. H. Wang, X. B. Wang, Y. Liu, F. C. Jiang and L. B.

21, no. 15, p. 4948, 2021.[4] X. L. He, Z. Q. Wang, D. H. Wang, X. B. Wang, Y. Liu, F. C. Jiang and L. B. Yuan,"Optical Fiber Sensor for Strain Monitoring of Metallic Device Produced by Very High- Power Ultrasonic Additive Manufacturing," IEEE Sensors Journal, vol. 19, no. 22, pp. 10680-10685, 2019.[5] A. Hehr, M. Norfolk, J. Sheridan, M. Davis, W. Leser, P. Leser and J. A. Newman, "Spatial Strain Sensing Using Embedded Fiber Optics," Additive Manufacturing of Composites and Complex Materials, vol. 71, no. 4, pp. 1528-1534, 2019.[6] S. K. Chilelli, J. J. Schomer and M. J. Dapino, "Detection of Crack Initiation and GrowthUsing Fiber Bragg Grating Sensors Embedded into Metal Structures through Ultrasonic Additive Manufacturing," Sensors, vol. 19, no. 22, p. 4917, 2019.[7] H. C. Hyer, D. C. Sweeney and C. M.

19, no. 22, p. 4917, 2019.[7] H. C. Hyer, D. C. Sweeney and C. M. Petrie, "Functional fiber-optic sensors embedded instainless steel components using ultrasonic additive manufacturing for distributed temperature and strain measurements," Additive Manufacturing, vol. 52, 2022. [8] D. Havermann, J. Mathew, W. N. MacPherson, R. R. J. Maier and D. P. Hand,"Temperature and Strain Measurements With Fiber," Journal of Lightwave Technology, vol. 33, no. 12, pp. 2474-2479, 2015.[9] J. Mathew, D. Havermann, D. Polyzos, W. N. MacPherson, D. P. Hand and R. R. J. Maier, "SS316 structure fabricated by selective laser melting and integrated with strain isolated optical fiber high temperature sensor," in International Conference on Optical Fibre Sensors , Curitiba, 2015.[10] M. Davis, J. Middendorf, N. Garg and O. J. O. III, "Additively manufactured componentswith embedded instrumentation," in ASME 2016 International Mechanical Engineering Congress and Exposition, Phoenix, 2016.[11] R.

Garg and O. J. O. III, "Additively manufactured componentswith embedded instrumentation," in ASME 2016 International Mechanical Engineering Congress and Exposition, Phoenix, 2016.[11] R. Zou, X. Liang, R. Cao, S. Li, A. To, P. Ohodnicki, M. Buric and K. Chen, "Optical fiber sensor-fused additive manufacturing and its applications in residual stress measurements," in Conference on Lasers and Electro-Optics, San Jose, 2017.[12] R. Zou, X. Liang, M. Wang, A. Yan, P. Ohodnicki, A. To and K. Chen, "EmbeddingDistributed Temperature and Strain Optical Fiber Sensors in Metal Components Using Additive Manufacturing," in Conference on Lasers and Electro-Optics, San Jose, 2018.[13] Y. Li, Z. Hua, F. Yan and P. Gang, "Metal coating of fiber Bragg grating and thetemperature sensing character," Optical Fiber Technology, vol. 15, no. 4, pp. 391-397, 2009.1079.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEmbedding fibre optical sensors into SLM partsPhilipp Stoll*1, Jinesh Mathew2, Adriaan Spierings1, Thomas Bauer1, Robert R. J. Maier2, Konrad Wegener31 inspire AG, St. Gallen, Switzerland 2 Heriot Watt University , School of Engineering & Physical Sciences, Edinburgh, UnitedKingdom3 IWF – Institute for Machine Tools and Manufacturing, ETH Zürich, Zürich, SwitzerlandCorresponding Author: stoll@inspire.ethz.ch, +41 71 274 73 27AbstractSelective Laser Melting (SLM) facilitates the integration of additional functionalities such as sensors into metallic parts. Such sensors can for example be embedded in sections of the parts non accessible after manufacturing. Additionally they can be positioned close to the region of interest.

Such sensors can for example be embedded in sections of the parts non accessible after manufacturing. Additionally they can be positioned close to the region of interest. Depending on the type of sensor it is even possible to monitor the structural health of the part itself. This paper discusses the integration of fibre optical sensors into SLM manufactured coupons in a commercial, industry scale, SLM machine. Such systems have merely limited accessibility for fibre handling. The embedment procedure as well as the corresponding bonding quality is explained in detail. Measurement results of the embedded sensor and limitations for the embedding procedure related to the use of commercial SLM systems are presented.

Measurement results of the embedded sensor and limitations for the embedding procedure related to the use of commercial SLM systems are presented. Additionally the need for further research is pointed out.Keywords: Fibre optical sensors, Embedding, Production environment, Commercial SLM system, Selective Laser Melting (SLM)IntroductionSelective Laser Melting (SLM) is an additive manufacturing (AM) technology that currently receives significant attention in the global production industry since it offers entirely new possibilities and strategies for the manufacturing of functional metallic parts. The prime benefits utilised are the option to produce highly complex internal and external geometries and the possibility to design a part based on its core functional purposes, not on the restrictions and limitations of the traditional manufacturing technologies.

This results in SLM parts typically exhibiting complex design features, bionic inspired shapes or topology optimized lightweight structures that are of enormous interest for various industrial sectors like medical engineering and aerospace. Another branch that starts exploiting the advantages that AM, and SLM in particular, have to offer, is the power generation industry since SLM facilitates not merely the manufacturing of high temperature loaded parts but also the integration of functional elements like sensors into these work pieces. Thereby the sensors can be placed in the immediate vicinity of the critical locations to be monitored. Thus the users of power generation units get a more detailed monitoring and consequently a more comprehensive understanding of the conditions within their systems, particularly during operation. The most promising variables to monitor the status of a system or a single work piece are temperature and strain.

The most promising variables to monitor the status of a system or a single work piece are temperature and strain. Knowledge of both parameters are essential to determine structural health and ageing conditions which are of crucial importance for a safe operation mode of highly dynamically and thermally loaded parts.A class of sensors that is well suited for measurements in high temperature applications are optical fibre sensors of similar or identical type widely used in telecommunications such as a 1815Corning SMF-28 fibre. There is a multitude of fibre sensor configurations available for a wide range of parameters with temperature and strain sensors being the most widely used as described by different authors, like Grattan [1], Majumder [2] and Li [3].

While Kashyap [4] explains that the sensing element can be for example a UV laser inscribed fibre Bragg grating which is a localised optical strain and temperature sensor, Islam [5] and Mathew [6] report on a Fabry-Perot type structure integrated into the fibre. Furthermore Srinivasan and Venkitesh specify in their book chapter [7] the usage of the fibre itself as a sensing element in which case the whole length of a fibre is the sensor and distributed temperature and strain date can be recorded along the fibre with high spatial resolution. Such fibre sensors are typically manufactured from fused silica material, which has a melting temperature of ~1720 °C and a strain point of typically 980 to 1050 °C.

Such fibre sensors are typically manufactured from fused silica material, which has a melting temperature of ~1720 °C and a strain point of typically 980 to 1050 °C. Hence, provided such fibres are handled and protected in an appropriate way they will survive a laser induced melt pool in a typical stainless steel material such as SS316 during a SLM process.Significant benefits of fibre sensors are their small size – typically 125 µm diameter – no electrical power requirements at the point of measurement, immunity to electromagnetic fields to name but a few. Fibre optic sensors employed in the project described here are based on wavelength encoded measurands which can be transmitted without loss of fidelity over long lengths of fibre.

Fibre optic sensors employed in the project described here are based on wavelength encoded measurands which can be transmitted without loss of fidelity over long lengths of fibre. If strain monitoring from within a structural component is required, Havermann [8] explains that the fibre requires to be bonded tightly to the component in order to transfer the component strain into the fibre, conversely, if temperature monitoring is required, there should be no strain transfer to the fibre as shown by Mathew [9]. Hence the exact mechanism and functionality determines the embedment strategy employed. In common with virtually all sensor technologies, fibre optic sensors are sensitive to temperature and to strain. This cross sensitivity has to be taken into account, thus – according to the patent of Willsch et al.

This cross sensitivity has to be taken into account, thus – according to the patent of Willsch et al. [10] – a strain sensor needs in most cases to be accompanied by a temperature sensor.The research presented in this paper deals with the embedding of optical fibres into 316 stainless steel coupons during the SLM manufacturing process on a commercially available machine. The procedure for the fibre integration is presented in detail focussing on the aim to achieve an appropriate bonding quality for sensing applications.State of the ArtThe integration of external elements into metallic parts during the manufacturing process has already been conducted by different research groups. The embedment of optical fibres is of particular interest since fibres facilitate the simultaneous measurement of two different values – temperature and strain – and induce merely tiny flaws or defects into the work piece due to their small cross sectional dimensions. Sandlin et al.

Sandlin et al. [11, 12] successfully embedded fibre Bragg gratings (FBG) in demonstrator parts of Inconel 600 using a vacuum brazing process. Despite the different thermal expansion coefficients between Inconel 600 and the fibre material promising results regarding temperature and strain measurements have been achieved. Kong and Soar [13] as well as Mou [14] and Norfolk [15] demonstrated that ultrasonic additive manufacturing (UAM) is another technology that facilitates the embedment of optical fibres in aluminium parts. In addition to the embedding procedure Schomer [16] reports in detail on the characterization of embedded optical fibres and thus shows the potential of integrated sensing elements and the feasibility of the UAM process to do so.Li [17] has shown in his dissertation for the first time that fibre embedment into metallic parts is feasible in powder based layered manufacturing.

This approach was based on a hybrid manufacturing process with alternating additive – laser metal deposition (LMD) – and 1816subtractive process steps. The results achieved by Li for the LMD technology inspired various researchers to think about fibre embedding during different other manufacturing processes. Recently this led to a significant increase in the number of research activities and publications dealing with integration of optical fibres in metallic parts during the SLM process. Maier [18] reports the first steps in the direct integration of metallic jacketed fibres into stainless steel coupons on a SLM test setup. In another paper [19] Maier emphasizes the aspect of developing optical fibres as sensors that are capable to survive the thermal loads during the direct SLM embedding process. Havermann [8, 20] focusses on the description of an embedment strategy for optical fibres into stainless steel coupons on a SLM test setup.

Havermann [8, 20] focusses on the description of an embedment strategy for optical fibres into stainless steel coupons on a SLM test setup. Furthermore he analyses the influences of various process parameters on the embedment procedure and the corresponding bonding qualities. In [21] Havermann shows promising results of strain measurements of integrated fibres. He even took these results a step further and applied the embedded fibres to measure the residual stresses in the SLM manufactured components themselves which allows a deeper insight into the SLM process and its characteristics [22]. Mathew conducted research on that topic as well, thereby focussing on the embedment of sensors for high temperature measurements up to 1000 °C [6].

Mathew conducted research on that topic as well, thereby focussing on the embedment of sensors for high temperature measurements up to 1000 °C [6]. All activities mentioned so far clearly reveal two challenges that need to be tackled by further experiment and investigations: all embedment trials have been conducted on a SLM test setup which allows the modification of many parameters and additionally offers excellent accessibility to the SLM process chamber which facilitates ease of handling of the fibres to be embedded and opportunities to tack down and hold in position. Whereas in commercial, industry scale, SLM system access is severely restricted, process interruptions are difficult to implement and fibre handling is awkward. The second issue is that the unidirectional application of the laser power results in a lack of bonding of the underside of the fibre.

The second issue is that the unidirectional application of the laser power results in a lack of bonding of the underside of the fibre. Although strain tests have shown that a single sided bonding is sufficient for 100% strain transfer, the remaining gap below the fibre is an issue which needs to be solved.Focus of the present studyThe research presented in this paper tackles the challenges described in the state of the art and thus focusses explicitly on fibre embedding on a commercially available SLM machine that leads to limitations and restrictions particularly with respect to accessibility compared to trials conducted on a SLM test setup.

Additionally the potential parameter modifications are considerably more restricted in a commercial SLM machine.Experimental ProcedurePreparation of the fibresIn order to ruggedise conventional optical fibres against the thermal processes of laser melting of high melt temperature metal powders surrounding the fibre, it requires to be protected by a thermal barrier in order to absorb the huge energy input of the laser during the SLM process and to provide a bonding layer between the fused silica fibre and the host material. The preparation of the optical fibres prior to the embedment procedure is currently a labour- intensive manual process that is of decisive importance for the subsequent process steps, however future automation should be able to address this issue. A cross section of the metallised fibre, showing the various layers that are vapour-deposited onto the silica core, is schematically illustrated in Fig. 1.

A cross section of the metallised fibre, showing the various layers that are vapour-deposited onto the silica core, is schematically illustrated in Fig. 1. The nickel coating facilitates an appropriate metallic bonding between the fibre and the surrounding coupon and acts as a thermal sink. The very thin chromium layer (<1 µm) acts as a keying layer between the fused silica and the electroplated Ni coating.1817Fig. 1: Cross section of coated optical fibre; left: metallic jacketed part of fibre; right: acrylate coated part of fibreSLM setup and embedding procedureThe experiments have been conducted on a Concept Laser M1 SLM machine with a process chamber of 160 x 230 mm. These dimensions lead to restricted accessibility to the manufacturing zone and limit the handling space, particularly for bending the fibres which however is inevitable since the SLM process requires a horizontal surface for the powder supply.In Fig.

2 the entire embedding process is illustrated in an exploded assembly drawing showing the single steps that need to be executed. The coupon that is equipped with an optical fibre is produced in 316 stainless steel and has a bounding box of 24 x 9 x 9 mm. Preliminary tests revealed that the cavity into which the fibre is mounted ideally has an elliptic shape with the semi-axes a = 570 µm and b = 200 µm (Fig. 3), if the outer diameter of the metallised fibre is 400 µm. The depth of the cavity – represented by semi-minor axis b – is of particular significance since it determines the thickness of powder that has to be molten after applying a new powder layer onto the tensioned fibre. If this layer is too thick – i.e. distance x in Fig. 3 is too large – sufficient bonding to the already consolidated material of the coupon cannot be guaranteed.Fig. 2: Embedding procedure, exploded assembly drawing1818Fig.

distance x in Fig. 3 is too large – sufficient bonding to the already consolidated material of the coupon cannot be guaranteed.Fig. 2: Embedding procedure, exploded assembly drawing1818Fig. 3: Form of cavity: elliptic shape with semi-axis a and b; metallised fibre located in cavityFurther preliminary tests indicated a positive effect on the surface roughness within the cavity if the top layer of the cavity is scanned a second time after its manufacturing applying a specific scan strategy reported by Yasa [23]. This effect of polishing the surface of the cavity is based on sufficient remelting and is realised with a small hatch distance in combination with low scan speed. The scan tracks are schematically illustrated in Fig. 2 and run parallel to the axis of the cavity, i.e. in the direction the fibre will be embedded. Fig. 4 shows a qualitative comparison between the surface roughness of an as-built by SLM and an additionally laser polished cavity.

in the direction the fibre will be embedded. Fig. 4 shows a qualitative comparison between the surface roughness of an as-built by SLM and an additionally laser polished cavity. The latter enhances the chances to achieve a decent bonding between the embedded fibre and the surrounding coupon, particularly on the underside of the fibre.Fig. 4: Effect of laser polishing on the surface roughness within the elliptic cavity measured with GelSight; left: unpolished, as built by SLM surface; right: laser polished surface, one additional scan with vectors parallel to axis of cavity.Tensioning of fibre The next steps in the embedding procedure are the tensioning of the fibre and the bonding by a specific scanning strategy. Since the most fundamental aspect is so assure that the fibre survives the interaction with the laser, the scan strategy has been developed prior to the tensioning concept.

Since the most fundamental aspect is so assure that the fibre survives the interaction with the laser, the scan strategy has been developed prior to the tensioning concept. Several different strategies have been tested using a nickel wire as a substitute for the optical fibres. All strategies first focussed on a fixation of the wire by spot- welding which is followed by the creation of a bonding layer in a second step. Fig. 5 illustrates a wobbling scan strategy to fuse the wire locally to the coupon and continue the SLM process on top of that layer. Although the fixation step proved successful, the further SLM build-up was not capable to properly fill the space between the wobbling scan tracks which resulted in voids and insufficient bonding between the first and the second part of the coupon. Based on that observation a bonding strategy that scans the entire coupon surface has been tested, eliminating the spot-wise scanning process step.

Based on that observation a bonding strategy that scans the entire coupon surface has been tested, eliminating the spot-wise scanning process step. The strategy that has been developed is divided into two scanning patterns that are shown in Fig. 6 as well as in Fig. 2, 1819named 4. perpendicular bonding and 5. parallel bonding. In the first step the scan tracks run perpendicular to the axis of the wire, in the second step parallel to it. Fig. 6 also demonstrates the way the wires are clamped on either side of the coupon – fixed clamping on one side and auto-adapting tensioning mechanism on the other. For the first scanning process the start point of the laser is of decisive importance to profit of the fibre mounting: it has to start at the fixed bearing and continue to the loose one enabling the wire to expand along the cavity but not to bend out of it in a consequence of the high energy input.

The thermal extension would cause the wire to bend out of the cavity if it was fix clamped on either side. A test has shown that the optical fibre behaves exactly in the same way (Fig. 7).Fig. 5: Wobbling scan strategy; left: schematically illustrated scan strategy; middle: fixed wire by wobbling scan strategy; right: cross sectional image showing insufficient bonding between the two parts of the coupon.Fig. 6: Bonding scan strategy; left: 1st step – scan vectors perpendicular to axis of wire. Start at fixed clamping; right: 2nd step – scan vectors parallel to axis of wire.Fig. 7: Fibre fix clamped on both sides bending out of cavity due to induced heat of SLM process.The precedent section has explained the importance of tensioning the fibre on one side. This is not trivial since the silica fibres are very brittle, particularly in the transition zone between the metallised and the acrylate coated sections.

This is not trivial since the silica fibres are very brittle, particularly in the transition zone between the metallised and the acrylate coated sections. Furthermore the SLM process requires a cleared horizontal surface for the recoating step which implies that the fibre has to be bent out of the processing plane. Although the bending radius of the fibre has been chosen as large as possible in compliance with the restricted space for handling due to the process chamber of the SLM machine, i.e. 14.2 mm, the bending generates an additional load on the brittle fibres that has to be taken care of in the process parameters. The mounting of the fibre has been realised with a cross bar and screws for the fixed clamping, while a spring provides a preload on the fibre and keeps the fibre tensioned during its expansion as a consequence of the laser energy input. Fig. 8 visualises the tensioning of the fibre.1820Fig.

powder layer before scanning - 3 layers: 30 µm each - No islands, scan vectors parallel to axis of cavity - 1. layer: 20 µm; scanned 1x - Standard parameter for 316 stainless steel - Islands strategy (5 x 5 mm)1821Bonding qualityThe bonding quality is analysed based on microscopic analysis of cross sectional images taken at different locations within the coupon. Since this requires a destruction of the embedded fibre, the evaluation of the bonding as shown in Fig. 9 has merely been performed on embedded nickel wire. The images clearly show two important results. While the bonding on the upper side of the wire has been realised without any pores or voids, there are still some issues regarding the bonding on the wire’s underside. However, these flaws don’t show a specific pattern – e.g. good bonding on one side of coupon, poor bonding on the other – which allows the conclusion that no systematic error is present.

However, these flaws don’t show a specific pattern – e.g. good bonding on one side of coupon, poor bonding on the other – which allows the conclusion that no systematic error is present. Consequently the embedment strategy can further be elaborated based on these results. However the bonding achieved would already be of satisfactory quality for temperature measurements.Fig. 9: Evaluation of bonding quality based on four cross sectional images per coupon. Top: coupon 1; middle: coupon 2; bottom: positions within coupons where different cross sectional images have been takenSensing capabilityThe goal of this paper is a detailed description of the embedment procedure of optical fibres into 316 stainless steel coupons which sets the baseline for further fibre integration experiment. To prove functionality of the embedded fibre the transmissivity for laser light through the fibre has been assessed. Fig. 10 shows both scheme and real setup of the measurement.

To prove functionality of the embedded fibre the transmissivity for laser light through the fibre has been assessed. Fig. 10 shows both scheme and real setup of the measurement. The result is presented in Fig. 11. The detector box records a current proportional to the incident light. The average of transmitted optical power through the metallised optical fibre prior to the embedment into the 316 stainless steel coupon is equivalent to 252.0 ±0.6 µA, while after the integration and embedment it decreased to an1822average of 242.5 ±1.0 µA (Fig. 11), which represents a typical loss of 3.8 %. A variation in transmission of a few percent is easily attributed to changes in connector losses. As has been mentioned in the introduction, the measurand is wavelength encoded hence the system is largely immune to a wide range of loss mechanisms at this level. Wavelength encoded sensors remain operational even if the total losses add up to 90% of total loss.

Wavelength encoded sensors remain operational even if the total losses add up to 90% of total loss. Thus, it can be deduced that the embedment procedure did not harm the optical fibre.Fig. 10: Light transmissivity measurement setup; left: schematically; right: real measuring setupFig. 11: Light transmissivity measurement through one fibre before and after its integration;mean ± standard deviation is shown for both conditions; several measurements have been conducted per condition to randomise the order and the influence of the connected elementsConclusions & OutlookThe experiment and results reported in the previous sections show that it is generally feasible to embed optical fibres into metallic parts on a commercially available SLM machine. This is the first step of bringing fibre integration from an experimental lab-environment to real production facilities. However, there are still some hurdles to be overcome.

This is the first step of bringing fibre integration from an experimental lab-environment to real production facilities. However, there are still some hurdles to be overcome. Besides the labour-intensive work to provide the nickel coating of the fibre in desired dimensions, which is not yet an industrially developed process but is conducted on an experimental setup, the handling of the brittle fibres as such as well as in combination with the restricted accessibility of the SLM machine during the embedment process has turned out to be the major shortcoming of the procedure. Furthermore the current embedding equipment and setup facilitates merely an integration of the fibre in a horizontally orientated, straight line, i.e. it is 1823not yet feasible to embed the fibre in a way that it follows the contour of a specific geometry. The shortcomings described above have to be taken care of in future trials.

it is 1823not yet feasible to embed the fibre in a way that it follows the contour of a specific geometry. The shortcomings described above have to be taken care of in future trials. However, first it has to be focused on repeatability of the experiment described in this paper. If those can be conducted with repeatable embedding quality and light transmissivity results, further analysis and measurements can be done. With regard to temperature measurements using the embedded fibres the bonding quality achieved so far and shown in Fig. 9 is sufficient, though for strain analysis within the surrounding material the fibre needs to be embedded without any flaws or pores which will be hardly possible to achieve with the approach shown in this paper. Another important conclusion of these experiment is that it is expected to be able to transfer the procedure to other materials, i.e.

bedding of Fiber Optic Sensors in Layered ManufacturingXiaochun Li, and Fritz PrinzRapid Prototyping Laboratory, Mechanical Engineering Department, Stanford University, CA 94305-3030AbstractLayered manufacturing enables integrating sensors during production of tooling or structural components. Sensors may be placed close to the points of interest prior to enclosure. Structures with embedded sensors are capable of monitoring parameters at critical locations not accessible to ordinary sensors. This work presents a methodology for embedding fiber optic sensors in metallic structures via Shape Deposition Manufacturing. Some of the main manufacturing issues are discussed. Embedded fiber optic sensors have been employed to measure temperature and strain. Proof-of-concept experiments on a non-contact fiber optic sensing system have been performed.

Embedded fiber optic sensors have been employed to measure temperature and strain. Proof-of-concept experiments on a non-contact fiber optic sensing system have been performed. By implementing the remote sensing system it is possible to measure temperature and strain of rotating components exposed to potentially hostile environments such as blades in gas turbine engines.IntroductionLayered manufacturing processes provide capabilities not easily achieved with conventional manufacturing [1]. Examples include the ability to create complex three- dimensional shapes with internal cooling channels, to produce parts with continuously varying material properties and parts with embedded mechanical and electrical components.

The possibilities also include embedding sensors and microprocessors to create "smart" products and embedding bearings, shafts and even bio-mimetic robotic mechanisms that have materials properties, tolerances, or finishes that would be difficult to obtain with traditional manufacturing process [2,3].Building parts in incremental layers allows complete access to the internal geometry of components. It allows for the placement of sensors close to points of interest and for their subsequent enclosure. Such embedded sensors can add intelligence and enable real-time health monitoring at some critical locations not accessible to conventional sensors. In addition, embedded sensors can also be protected from damage by extraneous environmental effects. Embedded sensors can be used to obtain information on the performance and structural integrity of functional components in service, which is of prime importance to many industries.

Embedded sensors can be used to obtain information on the performance and structural integrity of functional components in service, which is of prime importance to many industries. Examples include the manufacturing industry (molds, dies, drilling bits, etc.), the aerospace industry (components of jet engines), the oil industry (drilling equipment), the power industry (vessels and pipes), the automotive industry (components of motors), and construction industry (structural components in buildings). Past research has demonstrated the embedding of thin film sensors into metallic structures. In particular, embedded thin film strain gages have been314characterized in a four-point bending test and the results show good accuracy and linearity when compared with theoretical models and commercially available strain gages [4].Recently, fiber optic sensors and communication links were identified as promising candidates for integration into structural materials [5].

They offer immunity to electromagnetic interference, non-obtrusive embeddability, good resistance to hostile environments, and high bandwidth. One of the main problems for embedding thin film sensors into metallic structures is electrical insulation [4]. In contrast, fiber optic sensors do not need electrical insulation since they are not electrically conductive. Fiber optic sensors can provide sensing for structures operating at elevated temperature [6]. Furthermore, fiber optic sensors can be used to measure temperature and strain in a non-contact fashion [7]. This is especially significant when temperature and strain information need to be acquired from structures rotating in hostile environments, such as turbine blades. Traditionally, stationary flat or mildly curved plates are instrumented with surface mounted strain gauges and thermocouples [8,9]. For rotating blades, sensors are connected to the outside via slip rings.

For rotating blades, sensors are connected to the outside via slip rings. Naturally, such measurements yield low signal to noise ratios. In contrast, embedded fiber optic sensors can be employed for non-contact sensing of temperature and strain in rotating turbine blades.In this work, a manufacturing method is described for embedding fiber optic sensors via layered manufacturing. The characterization of the embedded sensors is presented. Moreover, a remote fiber optic sensing system is discussed which allows measuring temperature and strain in rotating structural components.Embedding Techniques for Fiber Optic SensorFiber optic sensors consist of modified optical fibers. Optical fibers are circular dielectric waveguides, as shown in Fig. 1, and have a central core surrounded by a concentric cladding with slightly lower (by 1%) refractive index. Fibers are typically made of silica with index- modifying dopants such as GeO2.

1, and have a central core surrounded by a concentric cladding with slightly lower (by 1%) refractive index. Fibers are typically made of silica with index- modifying dopants such as GeO2. A protective coating of cushioning material (such as acrylate) is used to reduce light loss due to microbending that occurs when fibers are pressed against rough surfaces and to provide protection to the fragile optical fiber core and cladding. The diameters of core and cladding for a single mode fiber are 9 m m and 125 m m respectively.The challenge for the embedding originates from the fact that the fiber needs to be embedded in high melting temperature metal structures. The fiber optic sensors will need to be protected during the high-temperature deposition steps. In particular, a protective layer is necessary to overcome temperature and stress experienced by the optical fibers during the formation of embedding layer.

In particular, a protective layer is necessary to overcome temperature and stress experienced by the optical fibers during the formation of embedding layer. Apparently, the soft and low temperature coating of the optical fiber must be replaced by a metallic coating. Electroplating of Nickel is selected. Nickel has high thermal conductivity, and bonds well to the subsequently deposited steel layer.The embedding sequence developed at the Rapid Prototyping Laboratory of Stanford University involve cleaning of fiber and substrate, sputtering thin film conductive layer, electroplating of Ni, and high temperature embedding. The process sequence is shown in Fig. 2.315CoatingCladdingCoreFig.1 Optical fiberLaser Deposited Stainless SteelNi electroplatingMetallic thin film coated optical fibersSubstrateFig.2. Explosive view of embedding sequenceCleaningThe substrates have to be cleaned before electroplating.

Explosive view of embedding sequenceCleaningThe substrates have to be cleaned before electroplating. To remove salts, oils and greasethe substrates are rinsed in borothene, acetone and isopropanol. Then the substrates are blown dry with dry nitrogen, rinsed with de-ionized water and dried again. The above steps are repeated once more and then the substrates are immersed in a 1:15 solution of micro-detergent in de- ionized water and are placed in an ultrasonic bath for 5 minutes. Finally the substrates are rinsed with copious amounts of de-ionized water and blown dry with dry nitrogen.The cleaning process of optical fibers is much simpler since optical fibers come with coatings. However, the polymer coatings need to be stripped before the optical fibers are embedded into metallic structures. Normally, a mechanical stripper is the standard tool for this purpose. However, the fiber optic sensors come in without coating in the sensing element.

Normally, a mechanical stripper is the standard tool for this purpose. However, the fiber optic sensors come in without coating in the sensing element. The interface between the cladding of sensing area and the neighboring polymer coating, as shown in Fig. 1, creates a stiffness mismatch and the fragile sensing element (without polymer coating) is easy to break during the mechanical stripping. More over, the mechanical stripping sometimes leaves residues of the polymer coating on the surface of the cladding. The residues may induce microbending and debonding during the later high temperature embedding. Thus, a chemical stripping process is applied. The optical fibers are immersed in a bath of acetone for about 10316minutes. The polymer coatings then detach the cladding of the fiber.

Thus, a chemical stripping process is applied. The optical fibers are immersed in a bath of acetone for about 10316minutes. The polymer coatings then detach the cladding of the fiber. The process yields no stress and no residues while the sensing elements remain undamaged.SputteringThe thin film deposition system consists of a stainless steel cylindrical chamber of 62l capacity (diameter: 46 cm, height: 38 cm), with two K.J. Lesker 1-KW magnetron guns. The bare optical fiber is sputter coated with a thin titanium film (< 1m m) to enhance adhesion. Then a thin nickel film (< 0.2 m m) is sputter coated over the thin titanium film. A short period (< 10min.) of Ni thin film sputtering is preferred since the Ni target becomes hotter and hotter during the sputtering. The resulting hotter plasma in the chamber may induce microcracks in the fiber optic sensor due to temperature gradients.ElectroplatingThe deposition of the nickel layer on the substrate is performed in two steps.

First the “seed” layer is grown in a nickel “strike” bath. Prior to the “seed” layer deposition the substrate is subjected to electropolishing by the application of a reverse voltage (substrate connected to the anode) for 2 minutes. Then the voltage is reinstated to its forward direction and the deposition commenced. Then the optical fiber can be placed on the substrate in a designated path (straight or curved), which can be created manually or patterned by machining or by LIGA/MEMS techniques. The other areas of the substrate are covered by insulating layers, which can be formed by photoresist, resin, or kapton tape. The substrate is then transferred immediately to the electroplating bath for the growth of the nickel layer up to about 0.5~1 mm thick. To ensure low residual stresses from the electroplating process, Barrett SN solution (Nickel sulfamate) plating bath is used. The temperature of the plating bath is controlled at 49(cid:176) C.

To ensure low residual stresses from the electroplating process, Barrett SN solution (Nickel sulfamate) plating bath is used. The temperature of the plating bath is controlled at 49(cid:176) C. A DC power supply is applied to control the current density. At the first hour of plating, the current density is set at 0.215 mA/mm2. Afterwards, the current density is raised as 0.646 mA/mm2. At the end of the deposition sequence, the substrate is retracted from the solution and all protective insulating layers (photoresist, resin, kapton tape) are removed. Finally, the substrate is rinsed in distilled water and blown dry.EmbeddingThe optical fiber can then be electroplated into thicker nickel layer or embedded to the other metallic structures by high temperature processing (casting, welding, arc spraying etc). In the Rapid Prototyping Laboratory, a 2400W, CW Nd:YAG laser is used to fuse metal powders.

In the Rapid Prototyping Laboratory, a 2400W, CW Nd:YAG laser is used to fuse metal powders. The powders are pre-placed on the surface of the nickel layer from a single powder feed nozzle. Laser power is controlled automatically and the deposition apparatus is moved across the surface of the substrate using a four degree-of-freedom robotic manipulator. By using this system, the composition of the deposit at any point on the surface can be accurately controlled.Process and Manufacturing IssuesThere are more issues involved. Sensors need to be placed at the points of interest inside the metallic components. We will focus our discussion on Shape Deposition Manufacturing (SDM) which was reported elsewhere [1]. The manufacturing cycle for embedding fiber optic sensors into functional parts can be depicted in Fig.

The manufacturing cycle for embedding fiber optic sensors into functional parts can be depicted in Fig. 3.317Part materialSupport materialDeposit (part)ShapeShapeDeposit (support)Insulating layerMetallic thin film coated fiber optic sensorInsulate and place sensorNickelShapeNi electroplatingFig.3. Shape Deposition Manufacturing cycle for embedding fiber optic sensorsThe main issues are the bond quality between the nickel layer and the substrate, thermal stress due to a large temperature gradient during layered deposition, and fragile ingress and outgress points of optical fibers from the metal part. During the manufacturing cycle, two types of defects, delamination and crack, will damage the optical sensors and the strength of the structure. The delamination or micro-crack can be caused by inadequate cleaning of optical fibers and substrates. Fig. 4a shows a crack in the boundary between an optical fiber and the electroplated nickel while Fig.

Fig. 4a shows a crack in the boundary between an optical fiber and the electroplated nickel while Fig. 4b display a delamination due to the inadequate cleaning of the substrate. Delamination or cracking may also originate during or after high temperature embedding. Laser cladding, which is similar to welding process, is used in this study. Cladding parameters for nickel are similar to those used to stainless steel, except the molten nickel weld pool is more difficult to control. The adequate mixing of the molten nickel and stainless steel power is crucial for good bonding. Fig. 4c shows a crack at the interface between laser-deposited stainless steel and electroplated nickel. In Fig. 4d, one can observe delamination, cracks and voids due to the inadequate cleaning, lack of molten metals mixing, and thermal stress.Several processing steps have been investigated to eliminate defects.

By optimizing the cleaning process and laser cladding parameters, the quality of embedding can be improved significantly, as shown in Fig. 5. The bonding strength between substrate and electroplated nickel can be enhanced by sand blasting the part before electroplating.318Electroplated nickel27 m mElectroplated nickel65 m mOptical fiberStainless steel substrate(a)(b)Laser deposited stainless steelLaser deposited stainless steelElectroplated nickel125 m mElectroplated nickel2 mm(c)Stainless steel substrate (d)Fig.4. Defects during manufacturing cycle125 m mLaser deposited SS316LNickelOptical fiber after electroplating(a)S.S304 (b)Fig.5. Embedded Optical fiber with good bondingSince SDM is an additive and subtractive process, the ingress and outgress points of the fiber optic sensors from the metallic structure need extra protection.

In particular, the optical fibers are very fragile during the “shape” process because chips from machining processes can easily cut off exposed optical fibers. One way to prevent this problem is to put nickel coating on the extra length, normally above 10 mm, of optical fibers outside the ingress and outgress points and to protect the extra fiber by using a polyethylene furcation tube with 1.2 mm diameter.319Another way is to completely eliminate the problem by cutting away all extra fiber leads outside the parts. This requires a non-contact sensing system to launch light into the optical fiber and to receive the light signal.

This requires a non-contact sensing system to launch light into the optical fiber and to receive the light signal. The new technique will be discussed later in this paper.Characterization of Embedded SensorsSince the sensor is coated with a nickel layer before embedded in stainless steel layers, it is necessary to study the characterization of the electroplated nickel as well as the sensor behavior under thermal and strain loads.The microstructure of electroplated nickel is so fine that Scanning Electron Microscope (SEM) is not enough to reveal its grain size. By using Transmission Electron Microscope (TEM), the microstructure of electroplated nickel is shown in Fig. 6. The grain size is about 150 nm. With such a fine grain, it is expected that the mechanical properties of nickel will be significantly improved. In a tensile test, the yield strength, tensile strength, and strain to failure are found to be 928.72 MPa, 946.65 MPa, and 4.6% respectively.

In a tensile test, the yield strength, tensile strength, and strain to failure are found to be 928.72 MPa, 946.65 MPa, and 4.6% respectively. Even without the high temperature embedding of stainless steel, the strong mechanical strength plus the corrosion resistance of electroplated nickel is suitable to produce embedded fiber optic sensors for special purposes, such as oil well and deep sea temperature/pressure monitoring.However, there is a thermal effect to the fine nickel microstructures during the laser cladding process. Actually the grain sizes become much larger than the electroplated ones. Obviously, the electroplated nickel recrystallizes during and after the laser cladding because the laser melts part of the nickel layer to mix it with molten stainless steel powder. Fig.

Obviously, the electroplated nickel recrystallizes during and after the laser cladding because the laser melts part of the nickel layer to mix it with molten stainless steel powder. Fig. 7 shows that the grain size of the nickel is decreasing from about 200 m m near the interface, between the laser deposited stainless steel (at right) and the nickel (at left), to about 10 um at a distance of 3 mm. The revealed microstructure changes are believed to result in lower mechanical strength but higher strain to failure.150 nmMicrostructures of nickel after embeddingLaser deposited Stainless Steel1200 m mFig.6. Microstructure of electroplated Ni (TEM)Fig.7. Microstructure of electroplated Ni after embedding320Sensor behaviorFiber optic sensors, such as Michelson, Fabry-Perot, and Bragg Grating, have been embedded in nickel and stainless steel structures. They provide higher sensitivity, good accuracy, and high temperature capacity.

They provide higher sensitivity, good accuracy, and high temperature capacity. For Michelson and Fabry-Perot sensors the potential operating temperature can be up to 1100 (cid:176) C (limited by the silica fiber itself) while for Fiber Bragg Gratings operation up to about 500 (cid:176) C (limited by the Bragg grating) is possible [7]. In this article, Fiber Bragg Grating (FBG) sensor is tol be discussed.FBG is compact, simple and can be demodulated in a wavelength-coding manner. When the FBG is expanded or compressed, it changes the grating spectral response. The Bragg condition is given byl 0 = 2 d ncore where l 0 is Bragg wavelength, d is the distance of pitches of the grating, and ncore is the refractive index of the core. The wavelength change due to temperature and strain is given by/l 0 = (a +z )D T + (1-pe)ewhere a is the coefficient of thermal expansion (CTE), z is the photothermal coefficient, and pe is the photoelastic coefficient.

By measuring the wavelength, temperature/strain can be determined. Fig. 8 presents the experimental results. The temperature sensitivity of the embedded FBG is almost two times higher than that of bare FBG. The strain dependence of the FBG matches the theoretical results very well [7].Rotating Parts with Embedded Sensors for Remote SensingSince light is used for the fiber optic sensors, it is possible to develop a remote sensing system to monitor the temperature and strain in rotating objects, such as turbine blades. The system is sketched in Fig. 9. After embedding the fiber optic sensors into a rotating part, a high speed wavelength tunable laser GCSR laser diode [10] can be used to launch light into the embedded sensor and the output signal, modulated by the temperature/strain, can be collected by a photo-detector or an optical spectrum analyzer (OSA).

Erbium-doped Fiber Amplifier (EDFA) is used to amplify the optical signal since light loss is a main concern due to the small size of optical fibers. Under a load (temperature or strain), the fiber Bragg grating reflects only single wavelength and all other lights with different wavelengths will go through to the photo-detector or OSA. Thus, at that single wavelength, the received light intensity will drop to almost zero. Thus, the Bragg wavelength under the load can be determined. The load can then be determined by the Bragg wavelength shift.To prove the concept of the system, a part with embedded FBGs is manufactured via SDM, as shown in Fig. 10. By applying thermal/strain loads, the wavelength shifts in the FBGs are monitored. Experimental results are shown in Fig. 11 with pictures scanned from the OSA. It clearly demonstrates the Bragg wavelength shifts after loads are applied. The result in light power domain (from photo-detector and oscilloscope) is shown in Fig. 12.

It clearly demonstrates the Bragg wavelength shifts after loads are applied. The result in light power domain (from photo-detector and oscilloscope) is shown in Fig. 12. The voltage is negative so that a pulse signal is shown instead of the proposed signal from the sampling scope in Fig. 9. All the results suggest the remote sensing system is feasible to monitor the temperature and strain variations for rotating turbine blades, if embedded with fiber optical sensors.321Fig.8. Experimental results for embedded FBGP(l )Embedded FBGGCSR laserEDFAFast tuning, 20nm/usP(l )OSAHigh speed sampling scopeP(l )Fig.9. Remote sensing system for rotating parts with embedded fiber optical sensorsFiber optic sensor pathMotorRotating partFig.10. Rotating part with embedded fiber optic sensor322Wavelengths from the laser diodeOriginal Bragg wavelengthNew Bragg wavelength after applying loadFig. 11.

Rotating part with embedded fiber optic sensor322Wavelengths from the laser diodeOriginal Bragg wavelengthNew Bragg wavelength after applying loadFig. 11. Bragg wavelength shifts on OSA for the rotating partReversed pulseFig.12. Pulse signal in oscilloscopeConclusionLayered manufacturing enables integrating sensors during production of tooling and structural components. Sensors may be placed close to points of interest prior to enclosure. This paper presents a method for embedding fiber optic sensors in metallic structures via Shape Deposition Manufacturing. Manufacturing issues, such as integrity of the structures with optical fiber sensors, are discussed. Embedded fiber optic sensors are characterized, showing linearity, high sensitivity, and accuracy. Proof-of-concept experiments on a non-contact fiber optic sensing system were proposed and tested.

Embedded fiber optic sensors are characterized, showing linearity, high sensitivity, and accuracy. Proof-of-concept experiments on a non-contact fiber optic sensing system were proposed and tested. Remote sensing is feasible to measure temperature and strain of rotating components exposed to hostile environments such as blades in gas turbine engines.AcknowledgmentThe authors are grateful to acknowledge the financial support for this work from Office of Naval Research under the contract of N00014-96-1-0354-P00003. The authors also like to thank Dr. G. Dehm from Max Planck Institute for Metal Research, Stugart, Germany, for323running TEM for the microstructure of electroplated nickel and Kapil Shrikhande of Optical Communication Research Laboratory of Stanford University for the valuable discussion and running the tunable lasers.References1. R. Merz, F. Prinz, K. Ramaswami, K. Terk, and L.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEmbedding of Liquids into Water Soluble Materials via Additive Manufacturing for Timed ReleaseCallie Zawaski1,2, Evan Margaretta2,3, Andre Stevenson2,4, Allison Pekkanen2,3, Abby Whittington2,4, Timothy Long2,3, Christopher B. Williams1,2 1Department of Mechanical Engineering 2Macromolecules Innovation Institute 3Department of Chemistry 4Department of Material Science and Engineering Virginia Tech, Blacksburg, VA 24061AbstractOne fundamental goal of personalized medicine is to provide tailored control of the dissolution rate for an oral dosage pill. Additive manufacturing of oral dose medicine allows for customized dis- solution by tailoring both geometric and printed material properties.

Additive manufacturing of oral dose medicine allows for customized dis- solution by tailoring both geometric and printed material properties. Direct processing of medicine via ﬁlament material extrusion is challenging because many active agents become inactive at the elevated temperatures found in the melt-based process. In this work, this limitation is circumvented by incorporating the active agents via in-situ embedding into a priori designed voids. This con- cept of embedding active ingredients into printed parts is demonstrated by the in-situ deposition of liquid ingredients into thin-walled, water soluble, printed structures.

This con- cept of embedding active ingredients into printed parts is demonstrated by the in-situ deposition of liquid ingredients into thin-walled, water soluble, printed structures. The authors demonstrate the ability to tune dissolution time by varying the thickness of the printed parts walls using this technique.Keywords: material extrusion, embedding, dissolvable materials, drug delivery1.IntroductionAdditive Manufacturing’s (AM) layerwise fabrication process has been shown to provide a cost-eﬀective manner for customizing products based on the speciﬁc needs of an individual. AM’s ability to customize parts has given the medical industry the advantage of being able to customize parts to a patients needs in areas such as in prosthetics, replacements, and tissue scaﬀolds. The ability for AM to provide custom, on-demand products has also been explored in the ﬁeld of personalized medicine.

The ability for AM to provide custom, on-demand products has also been explored in the ﬁeld of personalized medicine. AM possesses the ability to selectively control multi-material deposition on a voxel-by-voxel basis, allowing for more complex customization of oral dose medicine.To enable personalized medicine via AM, there have been many studies investigating diﬀerent matrix materials being processed by AM that successfully hold active agent [1–4]. Recent studies have demonstrated AMs ability to control the amount of drug inside an oral dose by either tailoring the concentration of drug or size of the pill [5–8]. The use of AM has also allowed for the fabrication of a single dosage tablet that incorporates multiple drugs [7–12].Another advantage AM can bring to the pharmaceutical industry is fast dissolving tech- nology and timed instantaneous release of dosage tablets.

For example, Spiritam uses binder2047jetting AM to create fast dissolving pills for treating epilepsy and myoclonic seizures. The fast dissolution is achieved by the powder process creating a porous structure for water pen- etration, coupled with a fast dissolving binder. The Food & Drug Administration (FDA) approved Spritam in August 2015 and is the ﬁrst 3D printed pill approved for oral con- sumption [13]. There has been signiﬁcant research in using binder jetting for fabricating drugs because of its ability to add active agents into the powder and/or binder [14–17]. This process is also preferable for its low operating temperatures, as operating temperatures 100 °C can damage active ingredients.In general, material extrusion AM uses higher temperatures than binder jetting, which limits its application for fabricating tablets for personalized medicine [1–5, 9, 18, 19].

To circumvent this limitation, the authors present a concept for in-situ embedding of active ingredients into a printed part to avoid exposing the active ingredients to the elevated pro- cessing temperatures. Melocchi and coauthors have demonstrated that material extrusion AM can create a liquid-tight part for fabricating a prototype capsule that could hold an adjustable dose [20]. While in Melocchi’s work the actives were added via a post-process, this work demonstrates that the actives can be added as a liquid in-situ.The goal of this work is to show how material extrusion can be used to embed active agents into a dissolvable material for controlled release. In this study, the process of em- bedding liquids is presented (Section 2) and is used to demonstrate that liquids can be successfully embedded during a print without leaking (Section 4). Tailored timed release is then demonstrated by adjusting the wall thickness of a printed capsule.

Tailored timed release is then demonstrated by adjusting the wall thickness of a printed capsule. This illustrates how a timed immediate release could be tailored to a patient’s application by adjusting the capsule geometry.2. Embedding liquids2.1. Embedding in AMEmbedding in AM is the process of building a part around preexisting components. Em- bedding is unique in that it takes advantage of inexpensive reliable previously manufactured components, and places it inside custom 3D geometry via AM. Meisel et al.

Em- bedding is unique in that it takes advantage of inexpensive reliable previously manufactured components, and places it inside custom 3D geometry via AM. Meisel et al. breaks down the embedding of parts during an AM build process into four general steps, Figure 1 [21].The part with a void corresponding to the embedded component is designed.The print is paused at the top layer of the void during the printing process.The embedded component is inserted into the designed void.The print is resumed on top of the embedded component.In this work, the general AM embedding process is extended to explore the concept of embedding liquids into printed parts. This process follows the general procedure described2048Figure 1: The general process for embedding as depicted by Meisel [21].by Meisel, with the exception that the print is not paused, and that the liquid is embedded in-situ. The reasons for this are described in Section 2.2 and Section 3.2.3.2.2.

The reasons for this are described in Section 2.2 and Section 3.2.3.2.2. Design considerations for liquid embeddingThere are several design considerations unique to embedding a liquid including the need for self-supporting structures, continuous printing, and accounting for shrinkage and thermal stresses.Self-supporting structures: Support structures are used for overhangs, bridging, and steep angles. Structures with enclosed voids, often need support material to support the roof of the enclosure. Typically a soluble material is used as a sacriﬁcial material for support structures; however, when parts are created using soluble material, dis- solving the support would also dissolve the part. Support structures can also be used without dissolving by creating brake away structures. Creating support structures that are manually removed adds an extra post processing step, and the removal process may cause enough damage to the part to break the liquid seal.

Creating support structures that are manually removed adds an extra post processing step, and the removal process may cause enough damage to the part to break the liquid seal. Therefore, designing part ge- ometries to feature self-supporting structures are important in creating a liquid-sealed part. This can be achieved by either using materials that can successfully bridge gaps or using the materials critical angle to enclose a void. Material bridging a gap requires that the material can hold an intended shape across the gap without a support struc- ture and properly bond between the roads creating the bridge structure. The critical angle can be used to angle inward to enclose a void. Structures with shallow angles (from the vertical) do not have enough material overhanging to necessitate a separate support structure. The maximum critical angle is dependent on the layer thickness, nozzle diameter, and material properties such as viscosity and surface tension.

The maximum critical angle is dependent on the layer thickness, nozzle diameter, and material properties such as viscosity and surface tension. The material properties are important to know how well the material can hold its shape while being deposited with some over-hang. 2049Liquid tight seal: Standard slicing software calculates machine code that has to dis- continue and then resume extruding numerous times throughout printing a complex part. This occurs when the print head stops at a designated location and moves to another location to resume extruding. At these start/stop points, the material does not bond well at the extruded end and the previously laid material. The weak bond introduces small gaps that create porosity in the part preventing the part from being liquid tight.

The weak bond introduces small gaps that create porosity in the part preventing the part from being liquid tight. A continuous printing loop, where the extruded material completes a full loop without a break in the extrusion, is needed to prevent leaks that can occur from start/stop defects.Preventing delamination: As the extruded material cools, the material solidiﬁes and shrinks. If a pause is used during the print, there is a weaker bond because there is less diﬀusion between layers [22]. As the layers shrink, internal stress is put on the part which causes parts to curl, or layers to delaminate. When layers delaminate, the liquid cannot be sealed. One way to prevent this issue is to print small parts where the thermal expansion is less signiﬁcant. Another way to prevent this issue is to not pause extruding and embed during the print (as noted above), so that there is not a weaker bond between layers.3.

Another way to prevent this issue is to not pause extruding and embed during the print (as noted above), so that there is not a weaker bond between layers.3. MethodsThe goal of this study is to ﬁrst explore if it is possible to embed a liquid into an encapsulated void processed by material extrusion using a water soluble polymer. The second goal is to explore how time of release of the embedded liquid during dissolution can be controlled by varying the thickness of a printed capsule.3.1. MaterialsA PVOH-based polymer, Mowiﬂex TC 253, made by Kuraray, was chosen for this study as it is able to be processed via traditional material extrusion and injection molding and is water soluble. For the purpose of directly printing tablets, using a material with a lower melting temperature such as the authors’ prior work in Poly(PEG8k-co-CaSIP) [23], would be better suited for the incorporation of low temperature active agents.

This work can use a material with a higher melting point because the embedded liquid, that could contain active agents, is not exposed to the high temperature of the extrusion nozzle. Mowiﬂex is adequate for the achieving the goals of this study: (i) demonstrating embedding a liquid with active ingredients inside a printed water-soluble container, and (ii) providing tailored dissolution through geometry.The material was characterized using Thermogravimetric Analysis (TGA) and Diﬀer- ential Scanning Calorimetry (DSC) to determine print parameters. The acceptable print temperature range is above the glass transition temperature, but below degradation temper- ature. TGA was performed on a TA Instruments TGA Q500 at 10 °C/min from 23 °C to 550 °C, with 1% degradation at 244 °C, as seen in Figure 2. TA Instruments Q2000 DSC was run using a heat/cool/heat cycle from -80 °C to 200 °C at a heating rate of 10 °C/min2050and quench cooled at 100 °C/min, Figure 3.

TA Instruments Q2000 DSC was run using a heat/cool/heat cycle from -80 °C to 200 °C at a heating rate of 10 °C/min2050and quench cooled at 100 °C/min, Figure 3. Data used from the second heating cycle, where the melting temperature was determined to be 177 °C.Figure 2: TA Instruments TGA Q500 curve for Mowiﬂex TC 253 at 10 °C/min from 23 °C to 550 °C.Figure 3: TA Instruments Q2000 DSC curve for Mowiﬂex TC 253 second heat cycle from -80 C to 200 C at a heating rate of 10 C/min.3.2. Experimental ProcessThe general process for this work is shown in Figure 4. First, the material was processed from raw material (1) into ﬁlament (2). The ﬁlament was then printed via AM material extrusion (3), in which liquid was embedded during the printing process (4).

First, the material was processed from raw material (1) into ﬁlament (2). The ﬁlament was then printed via AM material extrusion (3), in which liquid was embedded during the printing process (4). The ﬁnal part was subsequently dissolved in water (5), at which point the release of the embedded liquid was measured (6).2051Figure 4: A process diagram for 1-2) creating ﬁlament 3) ﬁlament material extrusion printing 4) embedding and 5-6) timed dissolution.3.2.1. Filament processingThe ﬁlament was made using a single screw desktop-scale extruder (Filabot Wee, Filabot, Barre VT). The processing temperature for creating the ﬁlament was at 198 °C, and run at a constant speed of 35 rpm. A 2 mm die was used along with a drawing system to draw the extruded ﬁlament to a target diameter of 1.8 mm. The slicing parameters for printing were adjusted to the actual ﬁlament diameter.3.2.2.

A 2 mm die was used along with a drawing system to draw the extruded ﬁlament to a target diameter of 1.8 mm. The slicing parameters for printing were adjusted to the actual ﬁlament diameter.3.2.2. Material extrusionThe material was printed on a custom delta-bot style material extrusion printer, driven by Marlin ﬁrmware. The printer uses an oﬀ the shelf hot end (E3D-v6, E3D), with the drive motor located close to the head for processing softer ﬁlaments. As informed by the DSC results Figure 3, the material was printed at 215 °C, with a bed temperature of 60 °C. The part was printed on blue painters tape to ensure adhesion to the bed. The part was cooled by forced convection to ensure that the material was below its melting point to hold its shape for the next layer.3.2.3. Liquid embeddingThe liquid was injected during the printing process via a syringe, which was driven by a pneumatically controlled dispensing unit (Ultimus V high precision dispenser, Nordson EFD).

Liquid embeddingThe liquid was injected during the printing process via a syringe, which was driven by a pneumatically controlled dispensing unit (Ultimus V high precision dispenser, Nordson EFD). Small amounts of liquid were injected every few layers, without pausing the print. The injections started once the print reached to the widest part of the part design. The2052liquid was injected so that the liquid was never touching the most recent 3-5 layers.This technique was developed through experimentation. Pausing during the print to embed the liquid tended to cause cracks in the geometry. The pausing caused a weak layer in the part, where the hot material was bonded to a cooled layer [22]. The diﬀerence in strength between layers caused the part to create a very small crack that was not always visible. These cracks typically appeared sometime after the part was created. A crack would take approximately an hour or two to form, causing the liquid to leak.

These cracks typically appeared sometime after the part was created. A crack would take approximately an hour or two to form, causing the liquid to leak. It is hypothesized that the crack in the layer was caused by residual stresses from the part shrinking and cooling. Another cause of cracks preventing a liquid seal was injecting all of the embedded liquid all at once. This caused the same problem because the room temperature liquid cooled the printed layers too quickly. Hence, why the liquid was then injected in intervals of small amounts during the print.3.3. Test Specimen Geometry DesignThe part was designed to take into account the design considerations from section 2.2 for material extrusion of a water tight part. Figure 5 shows the CAD design for the capsule. The part was designed with a small base to minimize any curling while also minimizing the embedded liquid heating from the bed.

Figure 5 shows the CAD design for the capsule. The part was designed with a small base to minimize any curling while also minimizing the embedded liquid heating from the bed. The part then angles outward (30°) to increase volume for the embedded liquid and creates a wider opening for embedding the liquid. The top of the part was closed in a pyramid shape (30°) to close oﬀ the part, sealing in the liquid. A pentagon shape was used to decrease the complexity and the ﬁle size of the part.Programmable release was demonstrated by varying the wall thickness of the printed part. The volume of the print material was kept constant, while the thickness of the wall was modiﬁed to be 1, 2, and 3 roads thick (0.4, 0.8, and 1.2 mm respectively). Figure 5a shows the road thickness direction. Table 1 shows images of the three diﬀerent wall thicknesses used for each of the parts and the calculated volume, cavity volume, and surface area.

Figure 5a shows the road thickness direction. Table 1 shows images of the three diﬀerent wall thicknesses used for each of the parts and the calculated volume, cavity volume, and surface area. The wall thickness normal to the angled wall was calculated using the angle. The volume of the material was kept constant for each wall thickness along with the aspect ratio of the gem shape. The size (height and width) of the gem was adjusted to keep the volume of the printed material constant, for the various wall thicknesses. As a result, the surface area of the parts varies in relation to the size of the part. The sample with the thickest wall had the smallest surface area. It would then be expected that the largest surface area would dissolve at a faster rate [24, 25].The part was designed using CAD (NX8.5, Siemens) and sliced using Slic3r v1.1.7 soft- ware. Some parameters in Slic3r were adjusted to ensure a complete liquid seal.

Some parameters in Slic3r were adjusted to ensure a complete liquid seal. An aligned seam was used between layers to keep the print continuous and prevented gaps created from starting and stopping extrusion. The layer thickness was set to be 0.2 mm. For the walls with multiple road widths, the perimeter was set to do the inner perimeter ﬁrst and the outer perimeter second. This and the small layer thickness helped to create a better connection between the layers at an angle.2053Figure 5: The capsule gem design for embedding a liquid here where a) road thickness is varied to change the surface area of the printed part and b) height is adjusted based on the wall thickness and keeping volume consistent.Table 1: Geometry measurements of the gem capsule design with diﬀerent wall thicknesses with a constant material volume. All measurements are from the CAD model.

All measurements are from the CAD model. The road thickness refers to the dimension shown in Figure 5a, and the wall thickness refers the thickness of the side walls from the outside surface to the inside cavity. The part bottom thickness was set to the road thickness.part and b) height is adjusted based on the wall thickness and keeping volume consistent.20543.4. Dissolution measurementThe part was dissolved in 200 mL water a 250 mL beaker with a stir bar at 400 rpm. The stir bar was oﬀset slightly to prevent any contact with the part. Images of the water against a white background were taken by a camera (GoPro Hero3) every 5 seconds to record the water color. The images were then processed via Matlab (MATLAB R2016a, The MathWorks, Natick, 2016). A consistent color square section in the image was ﬁrst selected. The square was then separated into its RBG values. The mean of the B (blue) values was then calculated and recorded as the value for that time stamp.

The square was then separated into its RBG values. The mean of the B (blue) values was then calculated and recorded as the value for that time stamp. As the blue liquid is released, the B value increases in the image value. The release is tracked by measuring the increase in the B value in the image sequence over time as the part dissolved and the embedded liquid is released.4. Results/DiscussionLiquid was successfully embedded during AM processing without the part leaking. Re- sults in Figure 6 show the measured embedded liquid release over time for each gem with diﬀerent wall thicknesses. As expected, the thinnest wall released the fastest. Figure 7 shows the average release time from a minimum of 3 samples for each wall thickness. The dissolution time was taken at the time where the embedded liquid was ﬁrst released. This graph shows that adjusting the wall thickness can control a time delay for an immediate release.

The dissolution time was taken at the time where the embedded liquid was ﬁrst released. This graph shows that adjusting the wall thickness can control a time delay for an immediate release. This is a similar concept as shown in Guptas work on programmable release. Gupta extruded PLGA solution and then directly injects emulsion ink into the aqueous core. The PLGA solutions contains diﬀerent lengths of AuNRs. The printed part released the emul- sion ink by using a laser to excite the nano-rods and ruptures the part [26]. In this study, programming release is done through a diﬀerent method using dissolution. Changing wall thicknesses creates a timed release, rather than using diﬀerent laser wavelengths to rupture the part.Comparing the release rate of each specimen, it is evident that the rate decreased as the wall thickness increased. This was caused by the loss of structural stability as the part took on water, causing the weight of the thicker walls to collapse on themselves.

This was caused by the loss of structural stability as the part took on water, causing the weight of the thicker walls to collapse on themselves. The collapsing walls trapped the liquid from immediate release and caused a slower release rate as the part continued to dissolve and release the embedded liquid. This is one reason for the larger standard deviation on the thicker wall release time. A way to avoid this problem is to have diﬀerent wall thicknesses throughout the part. So that a release wall would be the thinnest wall designed for the timed release and thicker walls could be designed for structural support.Other potential sources of error are in the ﬁlament and in calculation assumptions. The ﬁlament diameter had slight variations. The diameter used in slicing the ﬁle was the average which may have caused slight variation in the amount extruded during the print.

The ﬁlament diameter had slight variations. The diameter used in slicing the ﬁle was the average which may have caused slight variation in the amount extruded during the print. The calculations for the wall thickness are directly from the CAD model, but the gcode was created using Slic3r with 0% inﬁll, which may have caused slight variation in the actual wall thickness from the CAD calculated.2055Figure 6: The release of the embedded liquid for diﬀerent capsule wall thicknesses.Figure 7: The average release time for diﬀerent wall thicknesses. The release time is taken when the embedded liquid ﬁrst releases.20565. Conclusion & Future WorkThis work shows a process that enables embedding liquid into a printed part via material extrusion. Design considerations were established for self-supporting structures, liquid tight seal, and delamination.

Design considerations were established for self-supporting structures, liquid tight seal, and delamination. A process for embedding liquids was presented and demonstrated by the creation of water-soluble, liquid-containing vessels.The ability to tailor the time of release of an embedded liquid was demonstrated. The wall thickness of the enclosed void has a direct impact on time of immediate release of the embedded liquid. This could be used to create oral dose medicine that could create a burst release. Future work from the authors will explore geometries that include multiple chambers of embedded agents, where diﬀerent agents would be released at predetermined times.Other future work in this area includes creating the capsule out of drug-loaded material. Low melting temperature dissolvable polymers have recently been developed and can be processed via material extrusion [23].

Low melting temperature dissolvable polymers have recently been developed and can be processed via material extrusion [23]. A low-melting temperature polymer could be used to incorporate active agents in the capsule as well as an embedded material. Allowing for the release of multiple agents. Another area of interest would be embed powdered agents, where the powder could serve as a support structure to allow for more diﬀerent geometry creation. Embedding powder may also be more forgiving of materials and geometries that have diﬃculty processing a liquid tight part.6. AcknowledgementsThe author would like to thank the Long Lab and Whittington Lab for their collaborative work, Zane Haley for his work on with material properties, Dr.

AcknowledgementsThe author would like to thank the Long Lab and Whittington Lab for their collaborative work, Zane Haley for his work on with material properties, Dr. Staley and the Materials Science and Engineering Lab for access to use their machines, equipment, and training, and the Macromolecules Innovation Institute for the continued support and work in setting up collaborative projects, such as this study.References[1] M. Sadia, A. Sosnicka, B. Arafat, A. Isreb, W. Ahmed, A. Kelarakis, and M. A. Alhnan, “Adaptation of pharmaceutical excipients to FDM 3D printing for the fabrication of patient-tailored immediate release tablets,” International Journal of Pharmaceutics, vol. 513, no. 1-2, pp. 659–668, 2016.[2] T. C. Okwuosa, D. Stefaniak, B. Arafat, A. Isreb, and K.-w. Wan, “A Lower Tem- perature FDM 3D Printing for the Manufacture of Patient-Speciﬁc Immediate Release Tablets,” Pharmaceutical Research, pp. 2704–2712, 2016.[3] A. Melocchi, F. Parietti, A. Maroni, A.

bedding of SMD populated circuits into FDM printed objectsFlorens Wasserfall Department of Informatics, Group TAMS, University of Hamburg, Hamburg, GermanyAbstractThis paper introduces the concept of a highly integrated 3D-printing device which is capable of printing plastic parts with integrated, fully assembled electronic circuits in a single process. It is based on a standard FDM 3D-printer that has been augmented by a screw-driven conductive paste extruder for electronic circuit printing, a vacuum nozzle to pick and place SMD-components and a vision system to ﬁnd and precisely align the components before placing. To control the printer, an existing host software system has been extended to synchronize the communication with the printer for interactive operations and to generate the required movements from camera data by means of image processing. A number of objects, containing circuits on both the surface and inside of the object, has been successfully printed already.

A number of objects, containing circuits on both the surface and inside of the object, has been successfully printed already. Quality and durability of the generated parts have been evaluated by analyzing the curing characteristics of the conductive ink during the process and the adhesion of the components which are placed directly on the wet ink. The design concept aims for a practical, affordable approach that can be widely used by developers to lower the entrance barrier to the ﬁeld of 3D-printed electronics. Hence, the hardware is kept as simple as possible, avoiding complex and expensive components as laser or CNC-milling devices, focusing on algorithmic improve- ments in the preprocessing and control software. All developed hard- and software-components are available under open source licenses and compatible to common existing projects.1.

All developed hard- and software-components are available under open source licenses and compatible to common existing projects.1. IntroductionIntegrating electronic circuits has the potential to transform additively manufactured parts from pure mechanical objects into ”intelligent devices” by embedding sensors, actuators and controllers directly into the three-dimensional structure. A broad variety of approaches has been introduced over the last 15 years to address the integration of circuits into the Additive Manufacturing-process. The direct write approach has been successfully applied in combination with SLA fabrication to produce magnetic ﬂux sensors for the CubeSat project [3, 6, 10], LM555 timer based demonstrators [13, 10] and a highly integrated gaming die, containing a 3-axis accelerometer and several LEDs at the surface to indicate the result of a throw [11].

The focus is set to the advantages of three- dimensional component placement for applications that require a certain orientation e.g. to cover all three orthogonal dimensions of a magnetic ﬁeld. While the printing of conductive material is more or less automated, placing of components requires manual assistance. Several approaches have been reported to combine the FDM process with different conductive materials: integration of wires [1, 4], airbrush application of coating inks [14], deposition of several low melting metalalloys [15, 12], direct writing of conductive silver ink [4] and carbon ﬁlled plastic ﬁlament, called ”Carbomorph” [9].

Fundamental issues with the integration of external objects into the printing process have been considered from a general process-planning perspective [2] and evaluated for powder based printers [5].Based on this groundwork of research, the next consequent step is the integration of the entire process, including the assembling of electronics, into a single device. The FDM technology is considered to be well suited for integration of multiple materials and parts for several reasons: the surface is always ﬂat and open during the build process, no liquids or powders are covering the workspace, interruption and resuming of the print is possible without cleaning or registration of the workpiece. FDM technology is comparably cheap and widely available, increasing the num- ber of potential users. As stated in [4], printing conductive material directly on FDM-generated surfaces is subject to physical constraints.

As stated in [4], printing conductive material directly on FDM-generated surfaces is subject to physical constraints. The distortion caused by the surface roughness signif- icantly limits the achievable line width and increases the risk of broken connections. The authors propose a combination of additive FDM manufacturing and subtractive CNC milling to achieve a sufﬁcient quality. However, as demonstrated in this paper, the resolution is at least sufﬁcient for large package sizes such as the 1.27-pitch SOIC package used for the experiments, and can be potentially improved by aligning the print direction of the previous layer with the conductive trace.2. Description of the printerFigure 1 – First iteration of the printing system, based on a CNC milling device.The ﬁrst iteration of the introduced printer is based on a CNC cutter with modiﬁed driver software to gain compatibility with existing control software (ﬁg. 1). The printer can be controlled by any common host software.

1). The printer can be controlled by any common host software. It is equipped with a direct driven ﬁlament extruder, a syringe extruder for conductive paste processing and a pivot-mounted vac- uum nozzle to grip, rotate and place SMD- components. Prior to the printjob, the com- ponents are placed manually on a tray that is mounted at the same level as the printbed. The tray’s surface is divided into boxes by a raised rim. Each component is placed in a single box. This design enables developers to quickly create single prototypes without the need to obtain and handle large belt feed- ing systems. As described in section 4, the known box dimension is used by the vision system as a reference frame for self calibra- tion. Two cameras are attached to provide vi- sion for the object handling. One is mounted at the printhead, next to the extruders, facing downwards to locate the position of objects on the tray (top camera).

One is mounted at the printhead, next to the extruders, facing downwards to locate the position of objects on the tray (top camera). The second is mounted at the bed, facing upwards to correct the position and alignment of objects at the vacuum nozzle (bottom camera). All additional hardware parts are either widely available standard components or designed to be manufactured with a 3D-printer. The design is intended to be as modular and parametric as pos- sible to enable researchers and developers to reuse and adapt it to their requirements and existing hardware setup. The general procedure of printing a single plastic object including a functional circuit is depicted in ﬁg.

The general procedure of printing a single plastic object including a functional circuit is depicted in ﬁg. 2.Fabrication of a single plastic layerPick component from trayManual placing of SMD-components on the trayPlace component onthe surface or in amould directly intothe uncured inkTake picture and align rotation and offset of com-ponent at the gripperFinal ProductStart printjobDesign circuit(schematic)Merge designsDesign object(CAD)Direct wrinting of con-ductive paste intochannels at the surface layerFigure 2 – General process cycle for the fabrication of one object with integrated circuit. White ﬁelds indicate steps actually executed by the printer hard- ware, gray steps are pre- and postprocessing.3. Characteristics of the direct write extruderA screw driven syringe extruder is used to print the conductive material. The extruder consists of a gas-tight glass syringe, two motors, ball bearings a few screws and a number of printed parts.

The extruder consists of a gas-tight glass syringe, two motors, ball bearings a few screws and a number of printed parts. All components are comparably cheap and worldwide easy to obtain. Two conductive materials have been successfully printed with the extruder: In51Bi32Sn16 (Field’s metal) and silver ﬁlled polymer ink (#6130F, Methode Electronics Inc.). While In51Bi32Sn16 has a generally higher conductivity, a number of negative properties are prevailing. As stated in [15], the surface tension of liquid metal is very high, resulting in bulky traces and preventing the nozzle to disconnect from the ﬂoating material without strong oozing effects. Besides, the bonding be- tween the metal and the plastic object proved to be insufﬁcient. In previous work, lead based alloy has been extruded on PCL plastic by [8], but the bonding test was applied against a PCB surface only.In contrast, the #6130F polymer ink shows excellent bonding and extrusion characteristics.

The sheet resistance is 40 mΩ/(cid:3) as stated by the manufacturer. It was expected that the curing charac- teristics of ink completely enclosed in plastic might be insufﬁcient for more sophisticated circuits. To evaluate this issue, a number of test patterns was printed on an open surface and covered by plastic. The resistance of a 10 mm×0.5 mm×0.25 mm trace was measured over time at room tem- perature (20°C) and in a heated chamber (60°C). The results are shown in ﬁg. 3 and indicate that the curing is primarily determined by the temperature.Figure 3 – Curing characteristics of #6130F polymer ink. The resistance of a 10 mm×0.5 mm×0.25 mm trace printed on both the surface of an object and entirely enclosed in plastic is measured during the curing process at room temperature (20°C) and in a heated chamber (60°C).Depending on the nozzle, a line width of 0.7-0.2 mm is achievable by the extruder on a smooth surface.

A line width of 0.2-0.3 mm is desired for ﬁne pitch package sizes (e.g. TSOP). Currently, a line width of 0.5 mm can be reliably achieved on a printed surface. Smaller traces are subject to disruptions caused be bruises in the printed object. Besides, the line width is limited by the large amount of ink required to ensure adhesion of the parts as described in section 6.1.4. Vision aided pick and placeThe tray design described in section 2 requires active detection of parts for the gripping process. This is currently achieved by a system of two cameras. In principle, it would be sufﬁcient to use only one camera for both object localization and alignment, while the object lies on the tray, but the second camera signiﬁcantly improved the precision in our experiments.

The two main reasons for this effect are:the objects shift while the printhead moves the vacuum nozzle from the camera position to the part due to vibrations.The vacuum nozzle is not accurately rotational symmetric or at least looses its calibration after a few minutes of operation.In both cases, rotating the part results in massive offsets leading to misplaced components.4.1. PickSMD-parts must be placed on the tray by a human operator prior to a printjob. The rotation must be within a range of ± 45° since the vision system is not aware of the objects absolute rotation, e.g. the polarity of a diode. The tray boxes can be used as a reference frame to locate the objects, requiring only a rough calibration of the top camera. The rim of a certain box is extracted by running a Canny line detector (ﬁg. 4 left center) followed by a Hough transform (ﬁg.

The rim of a certain box is extracted by running a Canny line detector (ﬁg. 4 left center) followed by a Hough transform (ﬁg. 4 right center) constrained to a line length ofl >box size 2and a line orientation of θ =(cid:40)0° ± ε 90° ± ε.Cropping the image to the minimal bounding box of these resulting lines extracts a white back- ground square with a known edge length of box size. In a last step, the position of the SMD-part is obtained by computing the histogram and using the minima as the objects center of mass in relation to the tray coordinates (ﬁg. 4 right).Figure 4 – Image processing steps to locate a single SMD-object in a given tray box. From left to right: raw image provided by the top camera, edges extracted by Canny detector, relevant lines ﬁltered by Hough transform (green) and minimal bounding box (red), objects center of mass from histogram.4.2.

PlacePlacing an object requires accurate alignment, which can not be achieved by the human operator while preparing the parts on the tray. Again, Hough line detection is used to ﬁnd prominent features of the given object. A good threshold for the Hough algorithms accumulator can be determined by the expected size of the object from the extended gcode (section 5.2) or from the area covered by the object after separation from background. The Canny hysteresis threshold window is opened iteratively, starting with a high value, to adaptively ﬁnd only the most prominent edges (ﬁg. 5). In a second step the object is rotated to the desired placing orientation. This can induce a signiﬁcant deviation if the object is not exactly gripped at its center or if the vacuum nozzle is not exactly rotational symmetric (cheap hardware). Therefore, the objects’ center of mass is computed again from a second picture, taken after the rotation, to correct this effect.

Therefore, the objects’ center of mass is computed again from a second picture, taken after the rotation, to correct this effect. Implementing this in software allows the hardware to be quite imprecise and cheap while still achieving high accuracy. The part is then ”impressed” to its destination with a small negative Z-offset to ensure a strong adhesion.Figure 5 – Image processing steps to align an SMD-part and to correct offsets caused by prior processing steps. From left to right: raw image provided by the bed camera, main features of the object to determine orientation, object in target orientation (90°) and highlighted center of mass. The crosshairs indicate the vacuum nozzles pivot point.5. Host softwareConventional FDM-printers are usually operated by a realtime capable ﬁrmware, which directly controls motors, heaters etc. and a host software, running on a full-sized computer which inter- acts with the user and feeds the gcode to the ﬁrmware.

and a host software, running on a full-sized computer which inter- acts with the user and feeds the gcode to the ﬁrmware. This mainly requires unidirectional, linear communication. The extensions required to control the interactive pick-and-place system are im- plemented as a plugin for the Octoprint [7] host software. Both, Octoprint and the OctoPNP plugin are available under a GLPv3 license.5.1. Synchronizing the communicationComplex printing geometries are composed of a large number of small linear movements. The ﬁrmware usually buffers a number of commands to avoid interrupted movements caused by com- munication delays. Interacting with external objects during the printing process requires synchro- nization between the printers ﬁrmware and the host controller at some points to complete posi- tioning moves before taking a picture. This is realized by exploiting the M400 gcode command implemented by several ﬁrmwares.

This is realized by exploiting the M400 gcode command implemented by several ﬁrmwares. The ﬁrmware acknowledges every received command until the buffer is ﬁlled, and from this point on every time a command has been successfully executed. The M400 command blocks every acknowledgment messages for non immediately returning op- erations until it has been executed, causing the buffer to run dry.5.2. Gcode extensionsControlling the syringe extruder can be done by simply treating it as a normal extruder with high ﬁlament diameter (the syringe-piston diameter) and no heater. However, the vacuum nozzle for part handling relies on the data created interactively by the image processing, making it impossible to pre-compute the gcode commands. The proposed solution, currently implemented in the host software, is to extend the gcode for in-line part description as shown in ﬁg. 6. The XML statements are treated as comments and therefore ignored by a normal gcode interpreter.

1 ” y2=” −0.75 ” />Figure 6 – XML Gcode extension for in-line description of SMD parts. The M361 command actually ex- ecutes the placing operation for a certain object, during the printing process. The <shape> and <pads> information is currently used by the host software for visualization of the parts in the user interface. Inte- grating it as a threshold measure for the Hough algorithm is currently under development.6. Experimental resultsAs shown in ﬁg. 7, printing traces and precisely placing components on this traces works reliably on ﬂat surfaces. The orientation of the underlying layer signiﬁcantly affects the trace quality, causing the ”staircase” effect that can be noticed at the bot- tom of the green LED.6.1. Bonding characteristicsSince the parts are placed directly into uncured ink, the bonding characteristics are crucial to assess the electrical and mechanical quality of a manufac- tured object.

Bonding characteristicsSince the parts are placed directly into uncured ink, the bonding characteristics are crucial to assess the electrical and mechanical quality of a manufac- tured object. Besides, this information is relevant for the design of toolpath generators (slicing soft- ware). As shown in ﬁg. 8 (left), a series of specimen was produced on the printer, each containing several 1206 SMD-resistors, placed with increasing time delay on two 0.5 mm ink traces. After a curing pe- riod of one day, a shear force was applied to the parts by a calibrated ﬂat spring until the part was released. The results are shown in ﬁg. 8 (right). The data shows high variance, but a trend is clearly visible. The bonding quality partly depends on the conductive line geometry which varied over time due to several recalibrations of the syringe extruder during the long experiment.

The bonding quality partly depends on the conductive line geometry which varied over time due to several recalibrations of the syringe extruder during the long experiment. The results indicate that a fast placing of parts after the fabrication of a conductive trace is important to achieve reliable connections. A period of up toFigure 7 – Bonding and accuracy of SMD parts, correctly placed by the printer.SMD componentFCNC x-axisFlat springConductive inkFigure 8 – Evaluation of the bonding quality. Experiment setup to measure shear forces required to release SMD parts from cured ink traces on the surface of printed specimen (left) and results (right). The X-axis indicates the time between ﬁnishing the ink trace and placing a part into this trace.100 seconds for uncovered parts at the objects surface and 200-300 seconds for plastic covered parts is considered to be a reasonable value.

This implies that it is not possible to create a trace, continue printing the object and place the part into a left open cavity for most objects, because printing several layers usually requires more than 3 minutes. Potential approaches to mitigate this problem are:Embedding the parts upside down and connecting the part by ending the conductive trace on the pad.Deposition of small ink droplets on the pin landing areas. This requires an appropriate needle and potentially causes collisions between the plastic extruder and the object since the landing layer is below the objects surface.Utilization of conductive ink with slower curing characteristics (e.g. Methode #6130S).6.2. Printed objectsA number of objects was printed to evaluate different aspects of the behavior of conductive ink inside of FDM-generated objects.

Methode #6130S).6.2. Printed objectsA number of objects was printed to evaluate different aspects of the behavior of conductive ink inside of FDM-generated objects. Designing objects with integrated electronics currently requires extensive manual labor due to the lack of appropriate design software.Vertical connections Vertical connections between different layers have been successfully estab- lished by printing a wire with 45° inclination which is slightly distorted along the printing orientation to increase the contact area. A test connection printed at the surface for optical inspection is shown in ﬁg. 9 (left) and one of the microcontroller pins in the right image is connected to the objects bottom to establish a connection to the USB ports case.Surfaces The object in ﬁg. 9 (center) demonstrates the fabrication of free-form conductive sur-faces, e.g.

9 (center) demonstrates the fabrication of free-form conductive sur-faces, e.g. to be used as capacitive touch sensors.SMD populating To demonstrate the automatic placing of parts, a simple object was created, containing an ATtiny 85-20 microcontroller, controlling 2 LEDs (ﬁg. 9 right). The objectdirectly ﬁts into a USB port. The ATtiny was successfully programmed by a modiﬁed pro- gramming adapter connected to the USB pins, and the circuit is powered when plugged into a normal USB port.Figure 9 – Printed test objects to demonstrate vertical connections (left), capacitive free-form surfaces (center) and fully populated circuits (right).7. Conclusion and future workIn this paper, the design of a 3D-printer was successfully demonstrated which is capable of creat- ing complete electronic circuits integrated into printed plastic objects. This is achieved by enhanc- ing a common device with pick and place hardware, controlled by a computer vision system.

This is achieved by enhanc- ing a common device with pick and place hardware, controlled by a computer vision system. It shows that it possible to improve nowadays printing hardware for this task, at costs in the range of ∼1000 e . The design is modular and open, to be compatible to a broad range of hardware setups and applications. This is a basic step to enable researchers to further explore the potential of AM- manufactured functional objects.Future work includes the transformation of the hardware to ”standard” printers and the fabrication of more sophisticated objects. Two important aspects are the integration of sensors and the decen- tralization of electronics, superseding bulky circuit boards. A focus will be set to the development of design software, supporting the creation and pre-processing of objects.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEMBEDDING TEMPERATURE SENSORS WITH THE USE OF LASER-FOIL- PRINTING ADDITIVE MANUFACTURINGT. Turk1, M. Dougan1, L. Hendrix1, A. Reed1, C. E. Dominguez2, A. T. Sutton2, J. Park1, and M. C. Leu11Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409 2Los Alamos National Laboratory, Los Alamos, NM 87545AbstractLaser foil printing (LFP) is a metal additive manufacturing (AM) process using laser-based melting of metal foil technique. Using metal foils in AM has advantages over laser powder-feed processes, allowing for efficient heat transfer and resolving the drawbacks of powder-based AM such as potential powder inhalation health hazards, balling, spattering, and high powder costs.

In this study, we demonstrate the advantage of LFP for embedding sensors into structures using 304L stainless-steel foil as the feed material and two different types of temperature sensors. The first type is a resistance temperature detector (RTD) platinum sensor (Pt 100), and the second type is a k-type thermocouple probe. A detailed study of the sensor embedment through LFP revealed that the spot-welding scanning strategy significantly improves the product quality than conventional line-welding scanning strategy. As a result of this study, the feasibility of fabricating functional parts with embedded sensors using the LFP process is demonstrated.IntroductionAccording to the American Society for Testing and Materials (ASTM), AM is a general term for successfully joining materials layer-by-layer to form 3D parts, which is different from the traditional subtractive and formative manufacturing technologies [1].

There are a wide variety of AM techniques and categorization is based on the process architecture and characteristics. The following are the seven general AM categories [1]: binder jetting (BJT), directed energy deposition (DED), material extrusion (MEX), material jetting (MJT), powder bed fusion (PBF), sheet lamination (SHL), and vat photopolymerization (VPP). Two of these categories are similar to the process used in this study, laser foil printing (LFP), although it is not categorized in ASTM due to its unique processing technique. Sheet lamination (SHL) is a category that uses sheet/foil feed but fuses the layers using ultrasonic consolidation or adhesive joining. Directed energy deposition (DED) uses wire/powder feed, laser/electron beams, or electric arc to melt the feed material selectively. Laser foil printing (LFP) uses laser beams to melt the feed material of metal foils selectively.

Laser foil printing (LFP) uses laser beams to melt the feed material of metal foils selectively. Therefore, the method of LFP does not fall directly under one of the seven AM categories, but it is somewhere between sheet lamination (SHL) and directed energy deposition (DED).Since LFP was invented at Missouri University of Science and Technology [2], [3], some materials applications, characterization, and automation studies have been researched. As433examples, aluminum 1100 alloy [4], AISI 1010 low carbon steel alloy [3], 304L stainless steel alloy [5], [6], and zirconium-based amorphous metal [7]–[10] are material applications researched with the use of LFP. In addition, a comparative study between LFP and a commonly used laser powder bed fusion (LPBF) technology revealed that the finer grain structure and lower porosity of the laser-foil-printed part resulted in higher strength [6].

Moreover, the fast-cooling rate of LFP helps to form amorphous metal parts, which are difficult using other AM technologies and traditional manufacturing processes [7]–[10]. Most recent studies focused on automating the LFP system [11], [12] and improving process productivity with parameter optimization and laser polishing experiments [13].In the present study, we experimentally investigate the feasibility of sensor embedding using LFP additive manufacturing. The utilization of embedded sensors is gaining increased traction in both industrial and academic circles. To provide an example, the demands of Industry 4.0 encompass the integration of Internet of Things (IoT) technology featuring embedded sensors, applied across supply chain management, automated manufacturing, and complementary operational systems [14].

In addition, the sensors can be used for structural health monitoring via embedded sensors to obtain strain, residual stress, and temperature variations during the component’s lifecycle [15]. As a consequence of the layer-by-layer approach inherent in AM, these processes serve as more practical fabrication methods when compared to conventional manufacturing strategies for enclosing sensors into desired parts, such as capacitors, thermal and pressure sensors, resistors, microchips, structural electronics, and many others [16]. The design concepts of embedding the electronic components into metal parts for the LPBF process are presented in the research paper by Binder et al.[17], where they demonstrated that a weldable strain sensor could be embedded into a nickel alloy IN718 metal part by this process [18]. However, the use of a powder bed in the LPBF process makes the application of sensor embedding highly challenging.

However, the use of a powder bed in the LPBF process makes the application of sensor embedding highly challenging. Specifically, the process needs to be interrupted, the powder in the printed housing groove removed, and then the sensor placed during the embedding [19]. Moreover, process-related defects and anomalies such as keyholing, balling, and spattering are commonly known drawbacks of using powder-bed AM [20]. Additionally, Nunez et al. presented the embedment of thermocouples with DED [21]. The results showed that processing defects and porosities can be present, and the embedment strategy is critical. DED applications also include some hybrid processes using subtractive techniques[22] and ceramic deposition [23] to avoid defects to have improved surface quality and lower porosity. Yet, combining different processes would add additional steps and lower productivity.

Yet, combining different processes would add additional steps and lower productivity. Another application for embedding functional parts is Ramanathan’s research of piezoelectric polyvinylidene fluoride (PVDF) sensor embedding using ultrasonic additive manufacturing (UAM) [24]. UAM typically uses a milling operation to form the desired geometry and consolidates the sheets to form parts [25]. LFP uses a clean atmosphere without dust or unmelted powders in the formed grooves for embedding geometries, does not include any machining operation, and fully melts the foil feed resulting in dense metal parts with fine grain structure. On the other hand, LFP can also have high thermal loadings that can cause part defects and damage the embedded component during processing. In specific, foil welding has a lower spatter risk than powder welding. This is because the molten metal is less likely to be ejected from the weld pool when welding foil.

In specific, foil welding has a lower spatter risk than powder welding. This is because the molten metal is less likely to be ejected from the weld pool when welding foil. However, optimized scanning strategies can mitigate the drawback of potential processing issues. As a result, LFP can be a practical approach for embedding sensors and other types of electronic components into metallic parts.434The main objective of this study is to investigate the feasibility of embedding components into metal parts using LFP process. To showcase the potential of LFP for sensor integration with the capability of in situ measurements, temperature sensors have been selected as the components for demonstration purposes.

To showcase the potential of LFP for sensor integration with the capability of in situ measurements, temperature sensors have been selected as the components for demonstration purposes. This research aims to demonstrate how LFP can effectively embed sensors within metal parts during fabrication without damaging the sensor and the evidence of thermal decay.Overview of LFP process steps and systemThe LFP system consists of a continuous-wave (CW) infrared (IR) fiber laser for welding, and an ultraviolet (UV) pulsed laser for excess foil removal. The CW laser (IPG YLR-1000-MM- WC-Y11) subsystem includes a galvo mirror scanner (SCANLAB hurrySCAN 30) and an f-θ lens. The CW fiber laser has a center wavelength of 1070 nm, beam quality factor M2 of 3.04, and maximum average power output of 1000 W. The focal length of the F-θ lens is 330 mm, and the laser spot size is 160 μm. The UV pulsed laser (Coherent AVIA-355X) subsystem includes optical reflection mirrors and a focal lens.

The focal length of the F-θ lens is 330 mm, and the laser spot size is 160 μm. The UV pulsed laser (Coherent AVIA-355X) subsystem includes optical reflection mirrors and a focal lens. The UV laser's center wavelength is 355 nm, the pulse width is 30 ns, and the maximum average power output is 10 W. The focus length of the lens is 100 mm, and the laser spot size is 40 μm. Both the CW and UV laser beams are focused on the foil surface. In addition to the dual laser system, a 3-axis gantry system moves the specimen between the cutting laser and the welding laser. A circularly perforated foil clamping plate is used to apply pressure on the foil to hold the feed foil in place during the spot-welding step. A roller-to-roller foil supplier and a mechanical polisher are other system sub-components. Figure 1 shows the automated LFP system developed.Figure 1: In-house developed automated LFP system.

A roller-to-roller foil supplier and a mechanical polisher are other system sub-components. Figure 1 shows the automated LFP system developed.Figure 1: In-house developed automated LFP system. [11]LFP is a metal AM process to build a part via selective laser fusion of sheet feedstock material to a substrate. The LFP has multiple steps of (1) spot welding, (2) pattern welding, (3)435contour cutting, and (4) edge polishing all of which are schematically shown in Figure 2. Spot- welding step is the first step, and the foil clamping plate is used to compress the foil feed with the attached springs. Thus, spot-welding provides an attachment of the foil to the previous layer/substrate to prevent thermal distortion of the feed foil during the next step of pattern welding. In pattern welding, the material layer bonding occurs via the complete melting of the foil material. Then, contour cutting takes place to remove excess foil by detaching the layer from the feed.

In pattern welding, the material layer bonding occurs via the complete melting of the foil material. Then, contour cutting takes place to remove excess foil by detaching the layer from the feed. During UV cutting, burrs are produced; any non-flat surface features would result in defective welding for the following layer. Therefore, the final step is edge polishing to remove the burrs and flatten the part surface with a mechanical polisher consisting a #80 grit aluminum oxide grindstone.Figure 2: Illustration of the current process steps of LFP. From left to right: spot welding, pattern welding, contour cutting, and edge polishing. [11]Types of sensors in the embedding experimentsThe most common types of thermal sensors are thermocouples, mercury/alcohol thermometers, resistance thermometers, and radiation thermometers [26]. For demonstration we selected two different types of thermal sensors: thermocouples and resistance-based thermometers.

For demonstration we selected two different types of thermal sensors: thermocouples and resistance-based thermometers. Resistance thermometers, also known as resistance temperature detectors (RTDs), are made of pure material, such as platinum, to measure and calculate the temperature based on a linear relationship [26]. Thermocouples are composed of two metals to create a voltage between the hot sensor element and the cold reference junction [26]. Typically, thermocouples are less accurate than platinum RTDs, but the temperature range of measurement for thermocouples can be wider depending on the material type [26], [27]. The other considerations may include service life, the effect of pressure or vacuum, thermal decay or survival, and the thermal sensor cost in application [27].The advantage of LFP for embedding sensors into structures are demonstrated using 304L stainless-steel foil as the feed material due to its availability and ease of welding.

This foil purchased from Ulbrich Stainless Steels & Special Metals Inc. has a thickness of 0.005 inches (0.127 mm). Among the alternatives of temperature sensors available from the market, we selected two different types of sensors. The definitions, sensor types, dimensions, and temperature ranges for the two options are given in Table 1. The first sensor is an RTD, and it is selected because of its compactness and flatness facilitating embedment. It was purchased from the Innovative Sensor Technology IST AG with the product code 100513. The second sensor was selected because it has a weld pad, which allows LFP to perform laser welding onto the metal pad. It was purchased from Harold G. Schaevitz Industries LLC with the product code WPTC-ST-04-072-K-N.

It was purchased from Harold G. Schaevitz Industries LLC with the product code WPTC-ST-04-072-K-N. These two sensors with the cable and wiring assemblies are shown in Figure 3.436Table 1: Temperature sensor information#DefinitionSensor TypeDimensionsTemperature Range1Pt100, IEC 60751 F0.3 for high temperaturesRTD5 x 1.6 x 0.65 [mm]200 to +750 [°C]2WPTC weld pad surface temperature thermocouple probeK-Type25 x 25 x 0.25 [mm] (Pad dimensions)100 to +1300 [°C]Figure 3: Pt100 RTD sensor (left) and weld pad thermocouple (right).Spot welding as a scanning strategyIn the previous studies of LFP, the common scanning strategy of line raster scanning technique was used for the pattern welding step. Thus, the initial printing experiments used this approach. However, this scanning strategy needs to have a parameter investigation study to determine the best laser scanning parameters.

Thus, the initial printing experiments used this approach. However, this scanning strategy needs to have a parameter investigation study to determine the best laser scanning parameters. The results show that although the rectangular base plate could be printed successfully, printing a groove shape resulted in defects and unsuccessful printing, as shown in Figure 4. This is because the geometry of the groove shape has thin walls, which has a much slower cooling rate than the rectangular plate, leading to some heat accumulation and resulting in foil distortion and the Marangoni effect.

Therefore, a strategy of dense spot welding is investigated during the pattern welding step.437Figure 4: Sample printing results: (a) successful prints for a rectangular plate, (b-c) defective surface for the H-shape of Pt100 RTD, and (d) defective surface for the groove shape of WPTC.Spot scanning strategy is the application of laser welding discretely, which is different from continuous laser raster line scanning, see Figure 5. Calculations were done to determine the processing parameters for laser spot welding. According to Hung et al. [6], volumetric energy input (VEI) of the welding modes can be calculated with the equations 1 & 2, where P is the laser power [W], v is the laser scanning speed [mm/s], h is the hatch space [mm], s is the layer thickness [mm], t is the dwelling time [s], and d is the distance between spots [mm]. According to another research on the LFP process [5], the best laser welding parameters obtained had a VEI of 104.98 [J/mm3].

According to another research on the LFP process [5], the best laser welding parameters obtained had a VEI of 104.98 [J/mm3]. Thus, to achieve the same VEI, with 400 [W] laser power, 0.1 [mm] spot distance, 0.1 [mm] hatch space, and 0.127 [mm] layer thickness, laser dwelling time should be set to 0.333 [ms].Figure 5: Illustration of spot scanning (left) and line scanning strategies (right).𝑉𝐸𝐼𝑙𝑖𝑛𝑒 =𝑃 𝑣 ∙ ℎ ∙ 𝑠𝑉𝐸𝐼𝑠𝑝𝑜𝑡 =𝑃 ∙ 𝑡 𝑑 ∙ ℎ ∙ 𝑠The single energy input (SEI) for line and spot scanning are compared via theoretical calculations below. The length (l) of the scanned line varies between 1.6-24 [mm]. Thus, the following calculations include minimum and maximum single energy input for line scanning.

The length (l) of the scanned line varies between 1.6-24 [mm]. Thus, the following calculations include minimum and maximum single energy input for line scanning. The results show that the single energy input for the line scanning strategy is 16-240 times higher than the spot scanning respectively.𝑆𝐸𝐼𝑙𝑖𝑛𝑒 = 𝑃 ∙𝑙 𝑣438( 1 )( 2 )( 3 )𝑆𝐸𝐼𝑙𝑖𝑛𝑒−𝑚𝑖𝑛 = 400 𝑊 ∙𝑆𝐸𝐼𝑙𝑖𝑛𝑒−𝑚𝑎𝑥 = 400 𝑊 ∙1.6 𝑚𝑚 300 𝑚𝑚𝑠⁄ 24 𝑚𝑚 300 𝑚𝑚= 2.133 𝐽𝑠⁄= 32 𝐽𝑆𝐸𝐼𝑠𝑝𝑜𝑡 = 𝑃 ∙ 𝑡𝑆𝐸𝐼𝑠𝑝𝑜𝑡 = 400 𝑊 ∙ 0.333 𝑚𝑠 = 0.133 𝐽Figure 6 shows the dimensions for the geometries for the RTD sensor, weld-pad thermocouple, and the scanning path for the dense spots. The spot-welding pattern used in this study is a contour pattern that starts from the edges. In a previous research study, scanning inward mitigates the risk of foil distortion during processing [13]. The reason is that linear scanning in one direction may cause foil distortion due to the heat accumulation towards one edge.

The reason is that linear scanning in one direction may cause foil distortion due to the heat accumulation towards one edge. The red dots in Figure 6 are the approximate sensor measurement locations. The locations are essential for analyzing and investigating the collected temperature variation data.Figure 6: Dimensions of RTD sensor embedding geometry with sensor locations (left), contour pattern - spot welding path (center), and dimensions of WPTC sensor embedding geometry with sensor location.Pt100 RTD sensor embeddingThe laser-foil-printing experiments started with the computer-aided design of the Pt100 RTD sensor. First, the design files are prepared using SCANLAB laserDESK software, which is the laser processing software for the laser scanner used for welding. The sensor’s compact size allowed for embedding two sensors in one assembly.

The sensor’s compact size allowed for embedding two sensors in one assembly. Figure 7 shows the sensor and the double- sided design shape for embedding sensors.439( 4 )Figure 7: Image of Pt100 RTD sensor (left) and double-sided sensor embedding (right).Printing started with the base rectangular plate and then continued additional layer printing of the grooved H-shape until the depth was tall enough to embed the sensors. Afterwards, we manually placed the sensors into the groove. Finally, another layer is printed on top of the sensors to complete the sensor embedding. The pictures of the sample are given in Figure 8.Figure 8: Pt100 RTD embedding from left to right: grooved shape without sensors (first and second images), grooved shape with sensors, and after one layer printed on top of the sensors.Embedding temperature sensors helps to investigate the LFP processing in terms of in-situ welding monitoring.

The measurements were collected using a EXCELOG-6 Six Channel Handheld Temperature Data Logger for Thermocouples and RTDs. Figure 9 shows the temperature variation during the spot-scanning strategy. The oscillation is due to scanning being a contour path and the measured temperature increasing and decreasing accordingly. The scanning starts from the top edge, see Figure 6, where sensor 1 is closer to the melt pool, and the first peak in the sensor 1 measurement can be seen, then it cools down. Then the first peak in the sensor 2 measurement can be seen. This oscillation repeats until the entire surface is scanned.440Temperature vs.

Then the first peak in the sensor 2 measurement can be seen. This oscillation repeats until the entire surface is scanned.440Temperature vs. TimeRTD Sensor 1 (Channel 5)RTD Sensor 2 (Channel 6)45)C ° (40e r u t a r e p m e T3530253 71 15 19 13 27 21 35 39 33 47 41 5 9 5 5 5 Time (s)3 67 61 75 79 73 87 81 95 99 93 0 1Figure 9: The plot of embedded Pt100 RTD sensor readings during the LFP processing.Weld pad thermocouple (WPTC) embeddingAfter the study of embedding RTD sensors, embedding the second type of temperature sensor, i.e., the WPTC weld pad, was investigated. The embedding of this sensor started with its CAD model, which is shown in Figure 10 along with a photo of the sensor. Like the RTD sensors, the layer printing started with the base plate and continued with a groove shape with spot scanning.

Like the RTD sensors, the layer printing started with the base plate and continued with a groove shape with spot scanning. After the manual placement of the weld pad on top of the groove, the pad was welded on top of the surface of the previous layer. Then, the pad's exterior was cut using the UV cutting laser to maintain the rectangular geometry. The final layer printing on top of the sensor completed the embedding. Figure 11 and Figure 12 show the in-process LFP and sample pictures, respectively.

The final layer printing on top of the sensor completed the embedding. Figure 11 and Figure 12 show the in-process LFP and sample pictures, respectively. As seen in Figure 11, the processing is done in argon gas shielding atmosphere to avoid oxidation.Figure 10: Images of WPTC weld pad (left), embedded WPTC weld pad with translucent layers (center), and embedded WPTC weld pad (right).4417 0 1Figure 11: Laser foil printing of embedded WPTC: argon chamber assembly from top (left) and during spot welding (right).Figure 12: Weld-pad thermocouple embedding from left to right: printed base plate, grooved layers, welded pad after cutting, after printing another layer on top of the weld-pad.The printing of the layer above the embedded sensor has also been monitored for in-process temperature variations, and the result is plotted in Figure 13. The overall temperature rises due to the repetitive laser energy input in the plot.

The overall temperature rises due to the repetitive laser energy input in the plot. However, it oscillates during the input due to the sensor location shown in Figure 6. During the scanning of the edges, the weld spots are closer to the sensor measurement location, causing the localized peaks, and the cooling down causes the troughs when the laser spot welding is applied to the top and bottom edges.442Temperature vs. Time6560)C ° (e r u t a r e p m e T5550454035302505 10 35 40 65 70 95 0 10 2 15 3 10 5 15 6 10 8 15 0 5 9 1 2 1 2 2 Time (s)0 4 25 5 20 7 25 8 20 0 35 1 30 3 35 4 30 6 35 7 30 9 3Figure 13: The plot of embedded WPTC readings during the LFP processing.DiscussionIn this initial study, two different types of temperature sensors are used to investigate the feasibility of component embedding into metal parts using the Laser Foil Printing (LFP) process.

LFP has a clean fabrication environment and uses inexpensive metal foil as the feedstock, which is practical and cost-effective for sensor embedding purposes compared to popular powder bed metal AM techniques. This work showed that LFP is a viable means for embedding temperature sensors into metal parts. This can be further expanded to include strain and pressure sensors, electrical components, a wide range of different sensors, and many others.During the sensor embedding experiments, various process parameters for pattern welding using line scanning resulted in defects due to thin walls and heat accumulation. Monitoring and simulation techniques have been used to predict processing-related defects and anomalies in metal AM processes.

Monitoring and simulation techniques have been used to predict processing-related defects and anomalies in metal AM processes. To ensure an accurate prediction, researchers use various process monitoring techniques including the use of pyrometers, infrared (IR) and charge-coupled device (CCD) cameras, and non-destructive testing (NDT) (computed tomography - CT and x-ray) [28], [29]. Numerical simulations use mostly finite element analysis with commercial software packages [30] or graph-theory approaches [31] to investigate the temperature field. LFP is a relatively new process that has yet to be analyzed using monitoring and simulations. In the spot-welding scanning strategy in this study, the single energy input is too low to heat up the entire part surface, which mitigates the potential processing-related defects. Also, the single energy input does not vary with part geometry, which is highly desirable for simulating the process and predicting its outcome.

Also, the single energy input does not vary with part geometry, which is highly desirable for simulating the process and predicting its outcome. This scanning strategy has a high potential to avoid defects, but it needs further investigation into porosity, lack of fusion and/or delamination, mechanical properties, residual stresses, etc. in order to fully understand this technique compared with line scanning.4435 0 40 2 4Temperature readings can be collected using the embedded sensor for in-process monitoring of the process employed to embed the sensor. As a result, the monitored temperatures are significantly lower than the melting temperature of 304L stainless steel, which is around 1450 °C. This is because the volume of melt pool is small using the scanning strategy of dense spot welding, allowing the part not to be heated entirely. Moreover, any risk of heat accumulation- related defects (keyholing, Marangoni effect, and many others) is mitigated.

Moreover, any risk of heat accumulation- related defects (keyholing, Marangoni effect, and many others) is mitigated. Nevertheless, when positioning the sensors within the designated design geometry, there exist clearance gaps of 0.2 [mm] on each side. These gaps lead to reduced conductivity between the sensor and the embedding material (stainless steel), potentially leading to decreased accuracy in temperature readings. The conduction rate between the sensor and the metal part can be increased by filling the gaps with high-thermal-conductive adhesives.ConclusionsThis study investigates the embedding of two different types of temperature sensors using laser-foil-printing (LFP) additive manufacturing. The first sensor is a resistance temperature detector (RTD), and the second is a K-type thermocouple probe.

The first sensor is a resistance temperature detector (RTD), and the second is a K-type thermocouple probe. The process parameters in line scanning are crucial for printing successful parts with challenging geometries (e.g., thin walls) with embedding of sensors due to the high laser energy input. Therefore, an alternative scanning strategy is introduced in this paper: spot weld scanning. By using this scanning strategy, the successful embedding of these two temperature sensors into metal parts are demonstrated with thin-walled geometries using the LFP process.AcknowledgmentThis research work was partially supported by a research grant from the Los AlamosNational Laboratory.References[1][2][3][4][5]ISO/ASTM, “ISO/ASTM 52900: Additive manufacturing - General principles - Terminology,” International Standard, vol. 5, pp. 1–26, 2015, [Online].

ISSIVITY OF POWDER BEDSSamuel Sumin Sih and Joel W. Barlow Department of Chemical Engineering The University of Texas at Austin Austin, Texas 78712AbstractA simple model for estimating the emissivity of the surface of a powder bed from knowledge only of the bed porosity and solid emissivity is presented. Estimates by this model are compared with experimental measurements for powder beds of alumina, silicon carbide, and iron. Agreement within the uncertainty of the measurements, ± 10%, is obtained.IntroductionThermal models for the Selective Laser Sintering (SLS) process have been developed that can predict the sintering depth as a function of the energy received by the surface (1,2). The Energy Density, given byEnergy Density =p V*SCSphas been found useful for correlating sintering depth to SLS operating parameters such as laser beam power, P, beam scan velocity, V, and scan spacing, SCSP, for a particular powder that is processed in a particular geometry (2,3).

Equation (1) can not, however, be relied upon to predict the behavior of a new powder bed, nor can it predict fusion depth when the scan length is greatly altered. One of the reasons for this behavior is the use of beam power, P, instead of "power received", P', as heat by the powder bed. P' is related to P via,Its use requires knowledge of the emissivity, E, and absorptivity, U, of the powder bed, as well as the surface and ambient temperatures, Ts and TA , respectively, and is obviously more complex to implement. Part of this problem can be solved by the emissivity model presented and verified below.Predictive ModelWe begin with the usual assumption that most bodies are considered to be "gray" with no transmission of radiation. This leads to the assumption that U = E.

This leads to the assumption that U = E. The emission of radiation to ambient from a hot porous surface, such as the powder bed, can be assumed402(1)(2)to be caused by emission from the heated particles and emission from the cavities in the powder bed, all at the same temperature. Although the walls of the cavities are made by the same particles, the emissivity of the hole is a function of its geometry, particularly the fraction,/, of total cavity surface that is cut away by the emitting hole (4,5),b'H = -----=:....---- b's+f(l-b's)where b's is the emissivity of the solid particle and b'H is that of the hole. Equation (3) suggests that the hole emissivity is always higher than that of the material comprising the cavity and may even approach 1.0, the emissivity of a black body radiator, asf becomes small.

Equation (3) suggests that the hole emissivity is always higher than that of the material comprising the cavity and may even approach 1.0, the emissivity of a black body radiator, asf becomes small. Consequently, we expect that emissivity of the powder to be somewhat higher than that of the solid, and to follow the expression,where AH is the area fraction of the surface that is occupied by the radiation emitting holes. This expectation is experimentally met, see below.The general problem remains to describe the parameters AH andf in Equation (4) for a powder bed. German (6) provides a relationship that describes the mean pore diameter, d, as a function of the particle size, D, and the fractional packing density, p, for a powder bed comprised of randomly packed, single-sized spheres,d ='!:.. D ( 1- p)or3pd = 2 Drp 3 1- rpwhere <p is the fractional porosity of the bed, and <p = I-p.

D ( 1- p)or3pd = 2 Drp 3 1- rpwhere <p is the fractional porosity of the bed, and <p = I-p. German (7) also estimates that there are approximately 2.92 pores per sphere in the randomly packed bed. The surface of the randomly packed bed should have (2.92)2/3 or 2.043 pores per sphere in the plane.

The surface of the randomly packed bed should have (2.92)2/3 or 2.043 pores per sphere in the plane. The area fraction, AH, occupied by the pores is then,2AH=2.043 mi 4 + ;rD 422.043 mi 42Substitution of Equation (5) into Equation (6) yields,A H=0.9081.9082rp + 1403(3)(4)(5)(6)(7)We estimate/by considering each pore to be a cylinder oflength, L, so that12+ 4L dThe Length, L, can be determined from the definition of porosity, <p,by transposing and substituting Equation (5) to yield,L = O.514D l-lp lpSubstitution of Equations (10) and (5) into Equation (8) yields,and substitution of Equation (11) into Equation (3) leads to,5 s [2+3.082(I-lp)2]5lpH -5s [1+3.082(I-lpi ]+1lpEquation (12) suggests that the emissivity of the hole is independent of particle diameter and a function only of solid emissivity and the bed porosity.

As shown in Figure 1, the calculated emissivity of the hole decreases slightly as the bed porosity increases and also increases as the emissivity of the solid increases. The first trend occurs because/increases with increasing porosity, see Equation (11), causing the hole to become a less perfect radiator. The second trend simply notes the importance of the solid emissivity in determining the emissivity of the powder bed.404(8)(9)(10)(11)(12)0.950.90J= 0.85 ~ :> (j) (f) ~ 0.75 w0.800.700.650.60 ' - - - - -1 - - - - ' -_ - ' - - - - ' - _ J . - - - - 1 - - - - ' - _ - ' - - - - - ' - _ J . - - -A . - - - JOAOOA50.500.550.600.650.70FRACTIONAL POROSITYFigure 1. Calculated Hole EmissivityConlparison with Experinlental MeasurementsExperimental emissivities were obtained for three refractory powders, 15 ~lm Alumina (Norton), 7 pm Silicon Carbide (Norton/Lanxide), 4.6 ~lm Silicon Carbide (Norton/Lanxide),and <45 ~lm Iron (Hoeganaes).

These materials have measured bulk densities of 1.94, 1.54, 1.06, and 3.44 g/cm3, respectively, with corresponding fractional porosities of 0.512, 0.521, 0.670, and 0.563.The experimental apparatus has been described previously (8) and is shown schematically in Figure 2. The powder is contained in a thermally-guarded tube that heats the bed to the desired experitnental temperature. The powder is then to C02 laser radiation at its top surface.leiFigure 2. Schematic of Apparatus405A thermocouple is mounted 1 cm below the surface to record the temperature. Per experimental bed temperature, the temperature rise is limited to less than 10°C to permit thermal analysis that is based on constant thermal properties.

Per experimental bed temperature, the temperature rise is limited to less than 10°C to permit thermal analysis that is based on constant thermal properties. Surface absorbence, cr, assumed to be equal to surface emissivity, 8, is measured by noting differences in the thermal rise behavior of the powder bed between exposure, at the same laser power, of the natural surface and the same surface covered with a thin layer of carbon black powder, assumed to have t:R = a R =1.0. The thermal rise rate is measured in the linear region of the thermal rise curve in both measurements, and the bed heat capacity is assumed constant. Consequently, the governing equation ise=e(dT/ dt) R (dT / dt)R(13)Errors in this method are primarily those associated with determining the rise rates. We estimate relative errors in the range of ± 10%.Comparison between Equation (5) and experiment is shown in Figures 3-6 for the materials tested.

Comparison of Predicted and Measured Emissivity for 7 ,.un Silicon Carbide.TEMPERATURE. oK Figure 6. Comparison of Predicted and Measured Emissivity for Iron Powder.ConclusionsThe simple model for calculating the emissivity of a powder bed, presented here, seems to do an adequate job of predicting the effects of the emissivity of the solid and of the bed porosity. Obviously, more testing is necessary to prove the relationships presented, however these initial comparisons with experiment suggest that the model should provide good estimates that are accurate to within 10% by the proper consideration ofjust two parameters, the solid emissivity and the bed porosity. These relationships should prove useful for better understanding and controlling SLS operating parameters.References1. Nelson, 1. C.; Vail, N. K; Barlow, 1. W.; Beaman, 1. 1.; Bourell, D. L.; Marcus, H. L. Ind Eng. Chem. Res., 1995,34, 1641-1651.2. Nelson,1. c.; Xue, S.; Barlow, J. W.; Beaman, 1. 1.; Marcus, H.

PIRICAL MODELING AND VISION BASED CONTROL FOR LASER AIDED METAL DEPOSITION PROCESSMallikharjuna R. Boddu*, Srinivas Musti(cid:31), Robert G. Landers*, Sanjeev Agarwal(cid:31), and Frank W. Liou*Department of Mechanical and Aerospace Engineering and Engineering Mechanics (cid:31)Department of Electrical and Computer Engineering University of Missouri – Rolla, MO 65409AbstractThis paper gives a brief description of the laser aided manufacturing process. Empirical models describing the process dynamics of the laser aided metal deposition process is developed based on some of the models found in the literature. These models provide the basis for process planning and real time control. An embedded vision system, a two color temperature sensor, and a laser displacement sensor are incorporated for real time monitoring and control of the deposition process.

An embedded vision system, a two color temperature sensor, and a laser displacement sensor are incorporated for real time monitoring and control of the deposition process. The temperature profile of the surface and geometric characteristics of the melt pool are studied to ensure consistent operation of the process.IntroductionLasers have a tremendous impact in manufacturing industries. With laser innovations, the world in now experiencing the use of optical energy in a wide range of applications, from materials processing to Rapid Prototyping (RP). The use of lasers, in conjunction with metal powder, is one of the latest extensions to rapid prototyping, which had earlier involved plastic parts exclusively. Rapid prototyping using lasers has enabled the fabrication of complex, near–net shape functional metal parts directly from a CAD model at a low cost and offers faster turn around.

Rapid prototyping using lasers has enabled the fabrication of complex, near–net shape functional metal parts directly from a CAD model at a low cost and offers faster turn around. Currently this technology is implemented under a variety of names such as Direct Light Fabrication (DLF), Laser Metal Forming (LMF), Laser Engineered Net Shaping (LENS), Direct Metal Deposition (DMD), Selective Laser Cladding (SLC), etc. Though the system description and specifications of each of these vary, they all rely on the same principle of part fabrication, i.e. layer by layer deposition. A general description of the method of fabricating a part involves utilizing a laser to melt metal powder injected by a nozzle and laying down clad tracks via a positioning system having a controlled motion. In some cases the lasers may be directed along a defined path and tracks are laid down on a stationary table.

In some cases the lasers may be directed along a defined path and tracks are laid down on a stationary table. To control the deposition process it is necessary to understand the process system mechanics for which relations among various parameters need to be studied. The optimization of the process requires the measurements and control of parameters such as the powder feed rate, process speed, melt pool temperature and melt pool quality. During the last few years many sensors have been tried; the high cost of such equipment, the lack of real time or on-line control have led developers to seek other solutions. The use of laser displacement sensor, temperature sensor and CCD matrix camera with embedded image processing tools are sufficient to monitor all relevant process parameters.

The use of laser displacement sensor, temperature sensor and CCD matrix camera with embedded image processing tools are sufficient to monitor all relevant process parameters. These sensors with a standard data acquisition card and personal computer enable the operators to interact with the process and processing parameters.452Empirical Modeling of Laser Aided Metal Deposition Processes (LAMD)The process mechanics describing processes such as Laser Cladding, Laser Forming, Laser Ablation, Laser Welding, Laser Surface Treatment, Laser Metal Deposition etc., are the same, except for certain variations in the control parameters describing these processes. Some of these may include powder deposition, with or without material removal, or the enhancement of metal properties by heat treatment. These processes require certain factors such as dilution, heat- affected zone, porosity, clad dimensions, powder catchment efficiency, etc. to be monitored.

These processes require certain factors such as dilution, heat- affected zone, porosity, clad dimensions, powder catchment efficiency, etc. to be monitored. Identifying the control parameters and establishing relations to describe these factors forms the basis for empirical modeling. The first step involved is identification of parameters and categorizing them into dependent and independent variables based on certain assumptions. The need for these assumptions arises from the system’s flexibility to control certain parameters. Studies have to be conducted by varying these parameters and determining their effect on the part quality and system performance.

Studies have to be conducted by varying these parameters and determining their effect on the part quality and system performance. Figure 1 shows the classification and interaction of different system parameters (dependent and independent) describing the metal deposition processLaser beam Lens focal length di Nozzle stand off distance Beam mode structureParticle shape, size Powder thermal properties Nozzle geometry Carrier gas flowrate Shielding gas flowrate Type of feeder Nozzle stand off distanceSpot sizeLaser powerLaser power densityMelt pool t temperaturePowder flowrateDilutionClad widthClad heightTraverse speedOverlap factorPorosityCladding rateSurface finish(Misc. factors) Powder catchment efficiency Substrate temperature Interaction time Microstructure Layer bondingPRODUCT QUALITYFigure 1.

factors) Powder catchment efficiency Substrate temperature Interaction time Microstructure Layer bondingPRODUCT QUALITYFigure 1. General description of interacting parameters and their effect on quality of partThe second step involves developing analytical/empirical relations of the parameters describing the process. The final step involves fitting these relationships into models, performing specific experiments to evaluate the model constants, and specifying the conditions for which these relations hold.453A. Process Parameters Powder Feeder: Screw fed powder–feeding systems with a carrier gas is the general powder feeding system used in many deposition processes. The powder fills the screw either by weight or by the pressure of a carrier gas in the hopper.

The powder fills the screw either by weight or by the pressure of a carrier gas in the hopper. The amount of powder (in weight) being driven by a twin–fluted screw per revolution can be expressed asdρ2= rA where m is the powder mass per revolution (gm/rev), r is the feed screw helix lead (mm), ρ is the powder density (gm/mm3), and Ad is the cross sectional area (mm2) of one flute measured at a right angle to the screw axis. In practice, density depends on the hopper elevation, acceleration (if it is a mobile one), void factor (ratio of the space of air to that of solid), particle size, powder viscosity, and powder level in the hopper.

If Vm is the motor speed (rev/sec) then the powder mass flow rate M (gm/sec) can be expressed as (Li et al., 1993)mM =mmVBy controlling the amount of powder flowing into the melt pool, the energy input to the substrate can be controlled, thereby governing the amount of dilution that will occur (Blake et al., 1985).Energy Delivery System: The amount of heat input to a process depends on the system parameters. The heat delivered should be sufficient enough to melt the layers for proper bonding. Overheating results in more dilution and the part may melt back. The laser beam diameter at the laser material interaction is defined as spot size (D) and can be obtained from the relationFZDD b=where D is the laser spot size (mm), Db is the laser beam diameter (mm), Z is the laser nozzle stand off (mm), and F is the focal length of the focusing lens (mm).

Specific energy is defined asDvPE =where E is the specific energy (J/mm2), v is the traverse speed or feed rate (mm/sec), and P is the laser power (W).Powder Utilization Efficiency (ηηηηp): The powder utilization efficiency is expressed as the ratio of the powder delivered, to the powder deposited on the substrate within the duration of the laser irradiation on the substrate. Powder utilization efficiency also depends on material characteristics such as type, particle size, and shape of powder particles. It can be expressed mathematically as (Hu et al., 1998)p =ηWWcwhere ηp is the powder utilization efficiency, Wc is the net clad weight (gm), and W is the total weight of the powder delivered (gm). The total weight of the powder delivered is given bytW∫=0)( tMdtwhere t is the time (sec).454(1)(2)(3)(4)(5)(6)Lin et al.

The total weight of the powder delivered is given bytW∫=0)( tMdtwhere t is the time (sec).454(1)(2)(3)(4)(5)(6)Lin et al. (2000) performed experiments to determine the attenuation of optical energy through a focused powder stream and through a columnar powder stream for different nozzle stand off distances. This helps in determining powder utilization efficiency since this depends on the amount of energy being delivered to the substrate.Dilution: Dilution can be defined mathematically as the ratio of the clad depth in the substrate to the sum of clad height (above the substrate surface) and clad depth. It determines the amount of the liquid layer that needs to be formed on the substrate to ensure proper layer bonding. There are several parameters that determine the amount of dilution. These parameters include the powder flow rate, laser power delivered to the substrate, and table traverse speed. However the dilution mainly depends on specific energy.

These parameters include the powder flow rate, laser power delivered to the substrate, and table traverse speed. However the dilution mainly depends on specific energy. If specific energy is below a certain value (Kmin), bonding between the layers does not occur, i.e.KE <minorDvP<KminAlso, if specific energy is more than a certain value (Kmax), the previous tracks may melt back, i.e.KE >maxorDvP>KmaxThese values of specific energy can be used as indices in determining the range over which other parameters can be varied.Heat Affected Zone (HAZ): The heat–affected zone is the region beneath the bead which has a different microstructure as compared to the unaffected zone. This may result in cracking and surface distortions due to the hardening mechanism, particularly in steel. The width and depth of the HAZ can also be used as indices in determining the range of variation of other parameters.

The width and depth of the HAZ can also be used as indices in determining the range of variation of other parameters. Specific energy determines the depth of the heat–affected zone. For a given powder feedrate, a higher specific energy indicates more energy will be transferred to the substrate and the depth of the heat–affected zone will increase. Hu et al. (1998) performed experiments to show that penetration depth increases almost linearly with increase in the specific energy.Porosity: Porosity occurs due to cavities between tracks that form from overlapped tracks or due to the evolution of entrapped gases in the clad tracks. Porosity can be avoided if the aspect ratio (ratio of the clad width to clad height) is more than five and the percentage overlap is less than 70% (Steen et al., 1986).B. Process ModelThe quality and surface of finish of the part fabricated by the laser metal deposition process is determined by the clad geometry and its dimensional accuracy.

Process ModelThe quality and surface of finish of the part fabricated by the laser metal deposition process is determined by the clad geometry and its dimensional accuracy. Bead geometry can be defined by clad width and clad height, which depend on a number of other parameters. Empirical relations have to be developed to determine the effect of these parameters on clad geometry. Specific energy can be used to regulate the clad dimensions. For a given powder flow rate, the clad width can be regulated by varying the specific energy delivered to the substrate. Also, the specific energy depends on the laser spot size and traverse speed, which is evident from equation (4).455(7)(8)Williams and Deckard (1998) showed that the average part density at a particular specific energy is greater for a larger laser spot size. The range for which the traverse speed and laser spot size cause a variation in bead width needs to be determined experimentally.

The range for which the traverse speed and laser spot size cause a variation in bead width needs to be determined experimentally. Also, the range of traverse speed and spot size is limited by the amount of dilution as seen in equations (7) and (8). Thus, the relationship between bead width and, traverse speed and laser spot size is (Steen et al., 1986).DW =1(−av)where W is the bead width (mm) and a is an empirical constant. Hu et al. (1998) showed that an increase in powder flow rate results in wider cladding tracks provided the other variables are kept constant and there is sufficient energy available to form a clad. The net clad weight can be determined from the relationtWc∫=0)(η p)( tMtdtwhere Wc is the net clad weight (gm). Assuming the cross section of the clad to be a parabolic arch, clad height can be theoretically defined asH t3=2 WAwhere Ht is the theoretical clad height (mm) and A is cross section area of the clad (mm2). Hu et al.

Hu et al. (1998) performed experiments to show that the clad height increases with increased powder flow rate or specific energy, or with a decrease in traverse speed. An empirical relation can be developed to determine height, taking these factors in consideration.(12) where He is the experimental value of clad height (mm) and b is an empirical constant.

The experimental values of clad height can be compared with the theoretical values of clad height using equations (12) and (13).H e =bMvAspect Ratio: Aspect ratio is defined as the ratio of clad width to clad height and is expressed as HWA = (13)reFrom equations (10) and (13), the aspect ratio can be expressed asAr=D1(−av()bMv )/Cladding Rate: The cladding rate be expressed in terms of overlap factor (i), clad width (W), and traverse speed (v) as (Manjuluri, 2001) ) i Wv(15) 1( −= where C is the cladding rate (mm2/sec) and i is the overlap factor, which is defined as the ratio of the part of the bead overlapped to the overall bead width.C1( −=i1)(av)Dv−Sensors for LAMD ProcessThe quality of deposition process using laser aided metal deposition process is controlled by a large number of interrelated parameters. It is important to monitor and control these parameters to ensure consistent and efficient operation of the process.

It is important to monitor and control these parameters to ensure consistent and efficient operation of the process. We show in this article, that the use of a two-color temperature sensor, displacement sensor and a standard CCD camera enables us to control and optimize the laser cladding process.456(9)(10)(11)(14)The CCD camera is used to find the geometrical parameters such as the height and width of the track in real time. The camera also provides the temperature profile of the melt pool. This provides the operator with information related to process parameters like dilution factor and melt pool quality. Dual color infrared temperature sensors measure the surface temperature of objects without contact. The sensors work is based on the principle that infrared energy emitted by an object is proportional to its temperature. The temperature output signals can be input into a computer, for process monitoring and control.

Shape MeasurementsPreliminary experiments were conducted using CCD camera and imaging techniques toextract dimensions, thermal profile and quality of the molten pool in mild steel sample fabricated using the LAMD process. A total of 300 frames/run were recorded at a rate of 30 frames/sec with a CCD camera. The results shown are for mild steel material, at a laser power of 1500W and traverse speed of 10 mm s –1. In order to monitor the bead geometry during the process, the CCD camera is installed right on top of the nozzle head. The image captured in this angle enables us to compute the temperature variation in the melt pool. The images are captured during the off time of the pulsed mode operated laser. Simple algorithm is used to calculate the bead length and width in real-time. The bead height is obtained from the laser displacement sensor. This enables us to check the track dimensions being deposited online. The system is able to detect fluctuations in the mass feed rate.

The bead height is obtained from the laser displacement sensor. This enables us to check the track dimensions being deposited online. The system is able to detect fluctuations in the mass feed rate. As the mass feed rate increases, the height of the tracks also increases, up to a threshold, after which lack of power prevents the powder to correctly bond to the substrate. This threshold is a function of the laser power and the beam diameter. (Griffith et al., 1998) Figure 2 shows the digital image of the molten pool and the adjacent material created during the processing. The field of view encompasses approximately 4 mm.Figure 2. Sequence of images taken at a time interval of 1/60th of a second during laser off time457B. Temperature MeasurementsThe knowledge of the thermal behavior is critical to reliable and repeatable fabrication during LAMD processing. Our initial experiments show promise in both monitoring as well as understanding the thermal behavior.

Our initial experiments show promise in both monitoring as well as understanding the thermal behavior. During the cladding process the camera scans the melt pool and provides information of the temperature distribution within the melt pool. Of the images acquired, we take an average in order to remove unrepresentative features of the distribution like the movements within melt pool and presence of bursting bubbles during the process, which lead to large fluctuations in temperature. (Meriaudeau et al., 1997). By processing the images we obtain two gaussian distributions, the first one (low gray level) is related to the cold areas, the other gaussian centered on the average temperature contains information about temperature distribution within the melt pool. (Truchetet et al., 1996) The study of temperature profile of the melt pool provides us important parameters like dilution and melt pool quality.Dilution: The powder flow rate and laser power govern the dilution of the melt pool.

An uneven distribution of the temperature is the result of unmelted powder particles in the melt pool, due to lower laser power. This introduces porosity and poor mechanical properties. As the laser power increases, the surface temperature increases resulting in higher dilution. This causes the melting of previous tracks and residual stresses may be induced upon solidification (Steen et al., 1986).Melt pool quality: Thresholding is performed on the averaged image of the melt pool cross section to obtain the geometric characteristics and subsequently the melt pool quality. A well- formed melt pool has a definite boundary with finely distributed powder. An uneven or jagged boundary is a result of uneven distribution of the powder or high laser power.

A well- formed melt pool has a definite boundary with finely distributed powder. An uneven or jagged boundary is a result of uneven distribution of the powder or high laser power. This is one of the important aspects of quality control.Control Architecture for LAMD ProcessA closed loop control system is designed to monitor the bead width, bead height and melt pool temperature during the ongoing process and correspondingly the process parameters are adjusted in order to achieve desired performance. The control architecture to optimize the laser deposition process is shown in Figure 3. Adaptive optimal controller based on empirical process model is used to regulate laser power, feed rate, and powder flowrate to obtain an efficient cladding.Powder flowratePowder FeederAdaptive Process ControllerOutput SignalsLaser powerFeedrateLaserLaser Metal Deposition ProcessCNCData InputDisplacement SensorIR Temperature SensorImaging SensorImage ProcessingFigure 3.

Sensor Control Architecture458The two-color temperature sensor, displacement sensor and a standard CCD sensor, monitor the process in real-time and provide the process parameters (as shown in Table 1) to the adaptive controller. Adaptive controller provides control output in the form of powder flow rate, laser power and feed rate. The imaging sensor obtains sequence of images, which are processed using image processing tools to provide a feedback to adaptive controller. Empirical models, which describe the process dynamics, were used to design the adaptive process controller to optimize the LAMD process.ConclusionAn empirical model describing the process mechanics is presented for the LAMD process. A low cost system using laser displacement sensor, temperature sensor and CCD matrix camera has been presented which enables the operator to obtain information about laser cladding process. Some information is available in real-time and can be used in a closed loop control.

Some information is available in real-time and can be used in a closed loop control. The use of CCD technology enables determination of the clad width and temperature gradient of the melt pool. We showed that the data provided by our system enables us to improve the control of the cladding process through an adaptive process controller. Effects of some controller parameters on material processing are investigated. A statistically designed experimental matrix is used for this study.ReferencesBlake, A.G. and Eboo, G.M., 1985, “State of the Art Laser Hardfacing Using Dynamic Powder Feed Technology,” Conference on the Laser vs. the Electron Beam in Welding, Cutting and Surface Treatment, Reno, NV, pp.

pirical Similitude Method for the Functional Test with Rapid PrototypesUichung Cho Graduate StudentKristin Wood Associate ProfessorMechanical Engineering Department The University of Texas at AustinABSTRACTRapid prototyping has the potential to improve the performance of the design process both in cycle time and resources. Such improvements may be realized through the timely visual, ergonomic, and functional information provided by solid freeform fabrication (SFF) parts. Of these information classes, functional information is perhaps the least realized with current technology. A number of technical issues have limited functional testing of SFF parts, including sensor fusion, range of prototyping materials, part size etc. Our focus here concerns the material issues of functional testing, especially the potential differences in prototyping material choices to actual production materials.

Our focus here concerns the material issues of functional testing, especially the potential differences in prototyping material choices to actual production materials. For example, to derive accurate functional information of non-polymeric products from polymeric rapid prototypes, an improved similitude method that can overcome the distortion of material characteristics is necessary. In this paper, a new similitude method that utilizes specimen test data is introduced. This method develops a mathematical transformation between prototype and product behavior through specimen testing. This transformation replaces the role of the scale factor of the traditional similitude method, and provides a basis for relating prototypes to proposed production parts, even under dependent loading and material conditions.

Computational and experimental results of a structural design provide verification of the new method.1INTRODUCTIONDesign is an iterative process that recursively transforms customer needs into the appropriate set(s) of design parameters within various constraints, and the decision on design parameters are made from virtual and physical models. In order to maintain competitiveness, companies need to provide high-quality products at demanded time within allowable fmancial resources. Rapid prototypes provide effective visual, ergonomic, and functional information that is similar to the future product in a single physical entity, so it has the potential to accelerate the design process by providing required information with minimal time delay. The derived information is used to verifY virtual models or to correct design faults.

The derived information is used to verifY virtual models or to correct design faults. Among the various information classes, perhaps the functional information is the least realized.The required effort to build testable models is critical when the geometry of the target system is complex. Fortunately, a new technology, called rapid prototyping (RP) or solid freeform fabrication (SFF), enables effective fabrication of geometrically complex parts. Traditionally, a scaled physical model that can show similar behavior is designed on the basis of the Buckingham IT theorem, and the behavior is correlated to the target system behavior through the derived scaling laws [Kline 75, Baker 91].It is natural to expect various similitude studies with rapid prototypes, but very little literature exists on that subject [Dornfeld 95, Steinchen 95].

Some of the factors, that are limiting the utilization of rapid prototypes for the functional testing, are: (I) non-similar material characteristics of the rapid prototyping and production materials (e.g., nonlinear material properties); (2) distinct material structure (e.g., non homogeneous and anisotropic material properties of rapid prototypes); (3) limited material choices; and (4) restrictions on loading conditions (e.g., burning of polymer under combustion). Due to these problems, the559testing models fabricated from various rapid prototyping processes may not be suitable to make a precise prediction with traditional similitude methods.In order to extend the utilization of rapid prototypes for the reliable prediction of the functional behavior of future products, a new similitude method is developed and tested.

In comparison to traditional methods, where only the restricted information (the dimension of parameters and variables) is utilized, our new similitude method attempts to derive a local state transformation from the specimen pair(s) whose geometry is simple to fabricate. This novel specimen-based empirical similitude method is experimentally and computationally evaluated with three example problems. The new similitude method predicts the functional information well with one or two specimen pairs.2 Problems in Deriving Functional Information from Rapid PrototypesThere exist two main obstacles that lies in wide utilization of rapid prototypes for functional tests.

One originates from the characteristics of rapid prototypes, and another comes from the restrictions on the traditional similitude methods.2.1 Current Status of Rapid Prototyping TechniquesThe last decade has seen the emergence and continuous advancement of various rapid prototyping techniques, and at least 20 companies commercialized diverse rapid prototyping systems in Europe, Japan, and USA. Most of the systems can fabricate geometrically complex prototypes within 50 hours, where polYmers are the most popular base material [Aubin 94]. As rapid prototyping emphasizes the dramatically reduced fabrication cost and time, the loss or distortion of some information classes is inevitable.

As rapid prototyping emphasizes the dramatically reduced fabrication cost and time, the loss or distortion of some information classes is inevitable. The limited capacity (fabricable prototype size) and material choices of the rapid prototyping system are main factors that prevents full scale tests in various situations.2.2 Limitations of Traditional Similitude MethodsTraditional similitude methods, which are based on the Buckingham IT theorem, have been widely used in order to transform the behavior of physical prototypes to that of target systems [Kline 75, Baker 91]. The methods simply correlate (by multiplying scale factors) the corresponding behavior through a dimensionless parameter group. However, these methods have following limitations to make precise prediction of product behavior through rapid prototypes.(i) When two systems are governed by equations with different configurations, the prediction is erroneous.

The prediction of the deformation of a structure with small deflection from a structure with large deformation is the typical example of this type.(ii) The system parameters, that are related to the system behavior of interest, should be constant. However, the dependency of material parameters (e.g., thermal conductivity) on the system behavior (e.g., temperature) may not be negligible. In addition, the spatial distribution of material properties of the prototype and product are expected to be distinct, due to the unique fabrication schemes.(iii) All corresponding dimensionless parameters of two systems should be identical in the traditional methods.

Due to the limited material choices and difficulties in generating exact boundary conditions, the corresponding dimensionless parameters cannot always made to be identical.3 New Specimen-Based Empirical Similitude MethodThe above mentioned limitations oftraditional similitude methods seem to come from the utilization of restricted information, Le., the dimensions of parameters and variables. In comparison, our new similitude560method attempts to derive extended infonnation from the geometrically simple specimen pair(s) (one from the rapid prototyping and another from the production process) that can be easily fabricated.The new specimen-based empirical similitude method., that may overcome the limitations of traditional similitude methods, assumes that there exists a consistent local transformation between the corresponding n-dimensional sub-state vectors! of two systems only ifthe geometry is similar.

Based on this concept, the new similitude method derives a local state transformation matrix from the measured states of specimen pairs, and the derived state transformation matrix replaces the role of the scale factor in traditional the unknown states of the future-product are estimated from this derived state methods. transformation matrix and the measured states of the rapid prototype. In Figure I, the overall concept of the new similitude method is described., the state vectors are defined, and two plausible methods that realizes the concept are introduced in following subsections.Then,The state transition due to the material change is abstracted from the specimen pair.

As the state variation due to the geometrical variation is represented with the measured state vector of the target model, one can estimate the state vector ofthe target product.Model Specimen A geometrically simple physical entity fabricated from the rapid prototyping systemProduct Specimen A geometrically simple physical entityfabricated from the actual production processDerived transfonnation~_'f"''''''_Estimate the states of the target productTarget Model A geometrically complex part fabricated from the actual production process whose states are to he predicted(.

State measurement points)Target Product A geometrically complex part fabricated from the actual production process whose states are to he predictedFigure 1: Overall Concept ofthe New Similitude Method.Definition of State Vectors •2 ••• x tpd TSub-state vector ofthe target product, xtp = (x tp : k by 1 vector that unknown physical values of interest (e.g., temperature) on at k local points of target product Sub-state vector ofthe target model, x tm = (x tm : k by 1 vector that measured physical values of interest at k corresponding local points of the target model Sub-state vector of the i-th product specimen, x psi = (x P composed of measured physical values of interest at k corresponding local points of the product specimen Sub-state vector ofthe i-th model specimen, x msi = (xm...·! Xm.<i 2 '" x msi of measured physical values of interest at k corresponding local points of the model specimen! x tpis composed of! x tm2 ••• x tmis composed ofk ) Ttt •! x Psi2 ••• x psi k ) T): k by 1 vector thatisk ) T : k by 1 vector that is composedI The sub-state vector is defmed as n by 1 vector whose components are composed of measurable quantities to be predicted.

on a sub-domain Os in local coordinate space, where OscO and 0 is entire domain of the interest.5613.1 Similitude with Absolute Linear TransformationThe unknown relationship between x'P and xtm is approximated from the relation(s) of xpsi and xmsiAs an initial attempt, the relationship is assumed as a discrete linear transformation. Let's consider a linear transformation matrix T that satisfies following equationwhere T is a k by k matrix.In traditional similitude methods, the T is a diagonal matrix whose diagonal terms have identical state scale factor s. The scalar scale s is a real constant, and it is a function of N design parameters, DPi (e.g., characteristic length, material constants), and M loading parameters, p;-, as shown below.x. p s=-'- x. mIwhere Ri is a real number, and superscript p and m denote the product and model respectively.

p s=-'- x. mIwhere Ri is a real number, and superscript p and m denote the product and model respectively. Once all the DP;'sand F/s are decided, s becomes a constant real value, and the ratio of corresponding states of the product to model is same to the state scale factor s, at any point.In comparison, the new similitude method expects improved similitude results by empirically deriving the matrix T that satisfies equation (1).

The empirically derived transformation matrix T is generally not a diagonal matrix, and the elements, tfi'S, of the matrix T can be interpreted as the weighting factor on thej- th state of X",.d to predict the i-th state of xp • So, the state of the product specimen is not a scaling of only corresponding state of the model specimen but a balance ofthe considered neighboring states.d ·If states of n specimen pairs are measured, then the transformation matrix T should satisfy thefollowing equation,It is desired to minimize the number of specimens n, so the number of columns is usually lesser than that of rows, i.e., the size of the considered states. So, the number of unknowns are lesser than that of givenequations. When the [Xm.~l Xm.~2 ... xmsnr1does not exist, one way to find the T that satisfies equation (3)is to use the pseudo-inverse matrix [Strang 1988] as follows.T=[xPSl xps2 ••• xpsnJ.[xmsl Xm.~2 ... xmsnr-where A+ is the pseudo-inverse matrix of A.

xmsnr-where A+ is the pseudo-inverse matrix of A. Once the transformation matrix T is obtained, the state vector of the target product x 1p is derived from the following equation.x'P == T·x tmAs the transformation matrix T characterizes the state transition due to the material variation, the matrix for both specimen and target pair may be identical if the corresponding points' location in specimens and targets are well decided. If not specially stated, this approach is used as the new similitude method in Section 4.562(1)(2)(3)(4)(5)3.2 Relative Linear TransformationThe previous method assumes that the derived state transformation matrix T is independent of the geometry. However, the transformation may be dependent on the geometry.

However, the transformation may be dependent on the geometry. So, a method, that involves an alternative linear transformation ST that relates the deviation of states with respect to the reference specimen, is introduced as an adherent of the first method.In this method, the state vectors of the first specimen pair, xps1 and xms1 vectors. The variation ofthe state vectors of the second specimens are defined as, are considered as reference(6)where xps2 and x ms2 are state vectors of the second specimen pair. Similar to the previous method, the linear transformation oT that maps the state variation instead of the state itself.

is derived from the following equation.(7)Assuming that oT is constant (instead of T), i.e., the following equation is satisfied(8)x tp can be determined from(9)as xPS1, x tm and x ms1 are to be measured.4 EVALUATION of the NEW SIMILITUDE METHODSeveral experimental and numerical tests are carried out to test the validity of the new specimen basedempirical similitude method.4.1 Beam Experimentationcb~~~ljljl~t~~j~~1~1~lj~jlt~1~I1jljljlIml~1jlIjlj1j1jljljl~~jlm1jljlm~1FA1umiDum SpedJnen 1" X8" ( t=118")cPBPAluminum Target1" X8" (1=118")cFPGlycaroonale Spe<imeD2" X16" ( 1=1/4")PGlycarbonate Target2" X16" ( 1=1/4")(a) Specimen Pair(b) Target BeamsC: Clamping Point, F: Loading Point Figure 2: Aluminum and Polycarbonate Specimens and TargetsIn the beam experimentation, the deflection of the aluminum target beam at 1" (25.4 rom) and 7" In Figure 2, the geometry of the specimens and target(177.8 rom) from the clamping point is predicted.563beams are shown (excluding the clamping and loading zone).

The force on the polycarbonate specimen and target is scaled down to half of that applied to aluminum specimen and target, in order to emphasize the geometrical non-linearity due to the large deflection. For the convenience ofthe measurement, the LVDT is fixed to the same location without following the point of the beam when the large deflection is measured.Two possible sources of dissimilarity, that obstruct the reliable prediction of the deflection of the aluminum target with traditional similitude methods, are: (l) the nonlinear stress-strain characteristics of the polycarbonate; and (2) the geometrical non-linearity due to the large deflection.6(mm)6(mm)3,....----,.....-__-__.--_---r 2.5 ···········t············I············~·····)·/.·1-······ .; /i + . ! : / : ;.(1...... 1············/············1"···········t7(.~2: 1":.1.51 .•.•.•••.•.i.I...

! : / : ;.(1...... 1············/············1"···········t7(.~2: 1":.1.51 .•.•.•••.•.i.I... : ~ /! [0.5 ···../.·(.1l ~ '/'i ! [~ ~ 1'..F(N)023456023456(a) Vertical deflection at the near-clamping point(b) Vertical deflection at the far-clamping pointSolid line: Measured deflection, Dashed line: With the traditional method, Circle line: With the new method Figure 3: The Vertical Deflection of the Aluminum TargetThe vertical deflection of the aluminum target beam is predicted at two points (near-clamping point: I" (25.4 mm) from the clamping point, far-clamping point: T' (177.8 mm) from the clamping point) with the traditional and new similitude method.

As shown in Figure 3, the similitude result with the new method shows remarkable matching with the actually measured deflection of the aluminum target beam in comparison to the traditional similitude method.4.2 Computational EvaluationTwo [mite element (FE) simulations were performed with ANSYS to test the validity of the new similarity method in various situations. The assumed material properties for the FE simulations are as follows.4.0 -7.01000.0170.0 0.193Table 1: Material Properties for the FE Models4.2.1 + ~~========::::z'=/=====::::il ---+ • ~Static Structural ProblemII Vb=lOmm c::J b=2uun+ *=__=-"",r(__:::J"LoE_:::_"""::__::':I.-__::J ~ • ~bl=6mrn D hl=3mm 0 h2=2mmb2=4mm~L=lm (a) SpecimenL=lm(b) TargetFigure 4: Specimen and Target Beam564F(N)The problem is to predict the deflection of the aluminum target on the basis of the specimen testing results and the deflection of the polymer target.

The results on the angular displacement shows the similar results, even though it is not shown here due to the space limitation. In this simulation, the new method demonstrated its ability to map the behavior of two systems whose goveming equations take different forms. Besides, the new similitude method successfully predicts the behavior ofthe target beam when the Young's modulus is not constant.4.2.2Steady State Temperature and Thermal StressThe thermo-structural problem is introduced as an example to demonstrate the feasibility of the new similitude method when some of the dimensionless parameters are not kept to be identical. The geometry and boundary conditions of the considered specimen pairs and the target system are shown in Figure 6. The boundary condition at the outer surface of the polymer and aluminum specimens and targets are set to be same, assuming that one cannot easily control them (e.g., natural convection).

The boundary condition at the outer surface of the polymer and aluminum specimens and targets are set to be same, assuming that one cannot easily control them (e.g., natural convection). The difficulty in using polymeric rapid prototypes for the thermal similitude lies in the small thermal conductivity constant, and the low sintering temperature. Considering these material characteristics, the boundary conditions at the inner surface is determined as shown in Figure 6.b=1O(kW/m2 K). T=300Kb=IO(kW/m2 Kj, T=300KAdiabatic Surface(a) Specimen IAdiabatic SUrfacc (b) Specimen II(c) TargetFigure 6: Thermal Specimens and Targets565ll(m)In the two dimensional problems, the difficulty lies in deciding corresponding measurement points. In this example problem, characteristic points are defined as the point on the boundary surface with same boundary conditions, and the characteristic lines are defined as the line that connects two characteristic points.

Once the characteristic lines are determined, the lines are linearly divided to find out the corresponding measurement points. The steady state temperature are measured along the two characteristic lines as shown in Figure 6.As shown in Figure 7, the accuracy of the predicted temperature increases as the number of considered specimens increases. Even though the improved prediction accuracy is expected by considering more specimens, alternative method that utilizes the relative state transformation (see Subsection 3.2) has been applied and tested.

_:..295 ··············t·············t·············l..:·"·:~:··· ··t·············· . 290: ! . ~: . ! ~: ~ .. ~: :..~ .. ;...Figure 7: Similitude Results with theAbsolute TransformationFigure 8: Similitude Result with the Absoluteand Relative TransformationIn comparison to the similitude error in the temperature, the accuracy of the predictedthermal stresses (normal and shear) is remarkable as shown in Figure 9.<1.(NIm2) x 1n8 -109;......•..14.ot--I!t:8~..12..10,.28~.34!..5;..6.7r-.8!..5665 ConclusionsThe new specimen-based empirical similitude method predicts the various functional behavior (deformation, pure stress, temperature, and thennal stresses) of all considered example problems that cannot be accurately predicted with traditional similitude methods.

The new similitude methods with absolute and relative transfonnation have pros and cons their own, and a better transfonnation may be derived under the same concept of this paper - utilization of specimens to derive a consistent state transfonnation.Other problem domains (e.g., structural dynamics) will be tested to define the feasible application For the more realistic evaluation of the new similitude method, experimentation with SLSareas. prototypes will be carried out in the near future.6 AcknowledgmentsThe research reported in this document was made possible, in part, by a Young Investigator Award from the National Science Foundation. The authors also wish to acknowledge the support of Ford Motor Company, Texas Instruments, Desktop Manufacturing Corporation, and the UT June and Gene Gillis Endowed Faculty Fellow.

The authors also wish to acknowledge the support of Ford Motor Company, Texas Instruments, Desktop Manufacturing Corporation, and the UT June and Gene Gillis Endowed Faculty Fellow. Any opinions, findings, conclusions, or recommendations are those of the authors and do not necessarily reflect the views ofthe sponsors.References[Aubin 94] Aubin, R.F., "A World Wide Assessment of Rapid Prototyping Technologies," Solid Free/orm Fabrication Symposium Proceedings, Austin TX, August, 1994, pp.118-145 [Baker 92] Baker, W.E., Westine, P.S., Dodge, F.T., Similarity Methods and Engineering Dynamics: Theory and Practice of Scale Modeling, Elsvier, 1992 [Baker 67] Baker, W.E. and Westine, P.S., "Model Tests for Structural Response of Apollo Command Module to Water Impact," Journal o/Spacecraft, volA, n2, February 1967, pp.201-208 [Domfield 95] Domfield W.H., "Direct Dynamic Testing of Scaled Stereolithographic Models," Sound and Vibration, August 1995, pp.

12-17 [Elliott 69] Elliott, W., "Plastic Models for Structural Analysis," U S Naval Research Lab., Shock and Vibration Bulletin n40 pt 4, Dec. 1969, pp. 89-96 [Fay 93] Fay, RJ., "Scale Model Tests of Vehicle Motions," Vehicle and Occupant Kinematics: Simulation and Modeling SAE Special Publication n 975, SAE, pp51-54, 1993 [Jacobs 92] Jacobs, P. F., Rapid Prototyping and Manufacturing: Fundamentals of StereoLithography, Society of Manufacturing Engineers, McGraw-Hill Inc., 1992 [Kline 72] Kline, S.1., Similitude and Approximation Theory, McGraw-Hill Co., New York, 1975 [Steinchen 95] Steinchen, W., Kramer, B., Kupfer, G., "Photoelastic Investigation Using New STL-Resins," Solid Fre~rorm Fabrication Symposium Proceedings, Austin TX, August, 1995, pp.204-212 [Strang 88] Strang, G, Linear Algebra and Its Applications, Harcourt Brace Jovanovich, Inc., 1988567568.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAN EMPIRICAL STUDY LINKING ADDITIVE MANUFACTURING DESIGN PROCESS TO SUCCESS IN MANUFACTURABILITYPriyesh Mehta, Catherine G.P.Berdanier*, Manoj Malviya, Colin Miller and Guha Manogharan*Department of Mechanical Engineering, Pennsylvania State University, University Park, PA 16802Corresponding Author: cgb9@psu.eduAbstractThis paper characterizes engineering designers’ abilities to re-design a component for additive manufacturing, employing screen capture methods. Additive Manufacturing has garnered significant interest from a wide range of industries, academia and government stakeholders due to its potential to reform and disrupt traditional manufacturing processes.

The technology offers unprecedented design freedom and customization along with its ability to process novel and high strength alloys in promising lead times. To harness the maximum potential of this technology, designers are often tasked with creating new products or re-design existing portfolios of traditionally manufactured parts to achieve lightweight designs with better performance. To date, few studies explore the correspondence between design behaviors and manufacturability of final product within an Additive Manufacturing context. This paper presents empirical data from the design processes of six graduate student engineering designers as they re-design a traditionally designed part for additive manufacturing. Behaviors through the design task are compared between the study participants with a quantitative measure of the manufacturability and quality of each design.

Behaviors through the design task are compared between the study participants with a quantitative measure of the manufacturability and quality of each design. Results indicate opportunities for further research and best practices in design for Additive manufacturing and engineering education practitioners across multiple disciplines.Keywords: Design for AM, Topology Optimization, Design Framework, Eye-tracking, Manufacturability and Additive Manufacturing EducationIntroductionAn explicit focus on Design for Additive Manufacturing (DfAM) is growing to be an attractive avenue for researchers and industry to leverage the potential of Additive Manufacturing (AM), a field that has witnessed significant research and development in the last decade. While most research focuses on process and material properties, relatively few researchers explore the human contribution to the additive processes.

While most research focuses on process and material properties, relatively few researchers explore the human contribution to the additive processes. Recent DfAM research efforts seek to address this opportunity gap by developing novel frameworks to help accelerate implementation of design guidelines which would create components and assemblies best suited for AM [1-4] . These frameworks provide a skeleton for the iterative ideation and conceptualization process as a checklist to help designers create novel ideas for this technology. While a strict DfAM approach is inherently valuable, few researchers study how designers re-design an existing traditional part or assembly for AM.

While a strict DfAM approach is inherently valuable, few researchers study how designers re-design an existing traditional part or assembly for AM. The current trend in the industry for success in AM is to identify a set of66existing parts from product families designed for conventional manufacturing, and re-design it for AM [5].In such cases, there is a need for design guidelines, frameworks and workflows to help designers systematically approach the re-design process. Existing methods can be useful for application in the conceptualization stages of the re-design process, but designers still need assistance in making re-design decisions when looking to modify an established component[6]. Schmelzle et al provides a holistic framework approach with help of a case study towards re- designing an existing assembly by part consolidation [7].

Schmelzle et al provides a holistic framework approach with help of a case study towards re- designing an existing assembly by part consolidation [7]. An effort like the one adopted by Schmelzle directed towards developing a re-design workflow for shape optimization and weight consolidation is required for AM. While most DfAM research efforts aim to create a process/skeleton or framework to investigate the direction in which the re-design should be performed, there is a scarcity of literature studying explicitly how the re-design process occurs, especially with participants having basic or higher knowledge of AM theory and principles.This paper attempts to quantify and characterize the re-design methodology adopted by graduate engineering students when re-designing a mechanical component for the laser powder bed fusion (L-PBF) process using eye tracking and screen capture methods.

The structure of the paper is as follows: We provide a brief review or relevant literature, discuss methodological information for our empirical study including recruitment; participant profile; design prompt; data collection methods; and analysis protocols. The results explore the behavioral patterns of the designers through the re-design process and map them with a metric of quality for the final designs, evaluated using a proposed normalized manufacturability matrix for the L-PBF process. The findings motivate future research directions and implications for practitioners as discussed in the conclusion section.BackgroundDesign for Additive Manufacturing (DfAM) is the consolidation of shapes, sizes, geometric meso-structures, and material compositions and microstructures to make optimum use of capabilities of the AM process [8].

Organizations like General Electric and NASA have adopted DfAM approaches to achieve part consolidation and reduce the weight of the overall part without compromising the functionality of the part [9]. The development of knowledge of DfAM principles, rules, processes, tools and methodologies have been identified as one of the key challenges in mass adoption and implementation of AM, motivated by the realization that a designer’s lack of knowledge of AM principles prevent designers from optimally reaping the benefits of Additive technologies [10].Re-designing a traditionally-manufactured component for Additive Manufacturing can be performed using a variety of process and objective oriented frameworks, usually with human intuition and engineering decisions [1], combined with automated processes aided by design software tools.

One such tool is topology optimization (TO), a structural optimization tool used to optimize material distribution of structures to improve stiffness or other pre-defined objectives. Since the topology optimization algorithmic process removes material from all areas and locations where it is not required to support the specific loads or satisfy specific boundary conditions, the resulting geometries often contain structures that are not uniform in cross section. These structures sometimes resemble bones or tree branches; hence, the process is also known as bionic or organic optimization [11]. TO is a powerful approach for determination of optimum material distribution67under a specified design domain. However, there are several limitations associated with implementation of TO methods for AM; namely - mesh resolution, manufacturing constraints and post-optimization topology handling [12].

However, there are several limitations associated with implementation of TO methods for AM; namely - mesh resolution, manufacturing constraints and post-optimization topology handling [12]. Several research studies have tested the efficacy of designers with and without the use of proposed design heuristics and are proved to be effective towards achieving better and improved designs[13-16]. Re-design activities have also been classified into process-driven and designer-driven optimization, showing that a high possibility of the re-designed AM part becoming as much as 30 times more expensive to manufacture than the original design pressing in the need to validate the performance-cost tradeoff [17]. Further, there are frameworks associated with re-design methodology for AM focused on analyzing the end results of the design process, but to date no literature has been published characterizing engineering designers’ processes adopted while re-designing for AM frameworks.

This process is primarily driven by intuition and engineering judgements and hence it is worthwhile to investigate the cognitive process, spatial attention division and behavioral activity of designers involved in a re- design for AM task. Therefore, the research questions this study seeks to answer are as follows:1) What design behaviors do engineering designers employ when conducting a (re)Design for Additive Manufacturing task?2) How do designers’ behavioral patterns correspond with manufacturability metrics of the final designs?MethodsThe research design for this study employed human subjects research methods consistent with empirical research studies in design cognition and engineering education bodies of literature. After IRB approval, engineering graduate student participants were recruited to participate in a design challenge.

After IRB approval, engineering graduate student participants were recruited to participate in a design challenge. The following section outlines methodological considerations involved in recruitment, data collection, and data analysis.Recruitment and Participants. The participants for this research were recruited from a graduate level laboratory course at a large public university where the course objective was to provide hands-on experience with metal AM technologies. After obtaining IRB approval for the project, students were recruited to participate in a design challenge which involved re-designing a component for AM. Six master’s-level students chose to participate in the study. All participants had been enrolled in a specialized master’s curriculum in Additive Manufacturing and Design for at least one semester prior to data collection and hold at least a bachelor’s degree in an engineering discipline related to AM. Of the six participants, one was a woman.

Of the six participants, one was a woman. The number of woman participants, though low, are representative of graduate engineering populations in the United States [18]. The design challenge and data collection activities were conducted in the research team’s laboratory. Each participant conducted the design challenge individually on the laboratory machines, which are equipped with SolidWorks [19], Autodesk Fusion 360 [20] as well as eye tracking and screen capture data collection capabilities. The participants had been previously exposed to the workflow of laser powder bed fusion as part of the final class project. Therefore, it is assumed that the participants were aware of the preliminary opportunistic and restrictive design considerations along with post-processing workflow associated with the L-PBF technology.Design prompt: The part used for the re-design challenge was an airplane bearing bracket component from Alcoa Corporation.

This design was used as part of an open crowdsourcing competition by GrabCAD [21]. The goal is to minimize the mass and optimize for weight and68strength while fitting within the target envelope and meeting the technical requirements. The design prompt presented to the participants included re-design objectives, design requirements, loading conditions and material properties for simulating performance. The prompt also indicated that the part had to be re-designed for the laser powder bed fusion technology as shown in Appendix A. The intent behind selecting an open source design was to benchmark against a well- studied case for AM re-design which allowed the researchers to focus more on the re-design activity and decision-making process rather than investing time in creating a new design with loading conditions. There are other advantages of using open source competition such as cost, sustainability and quality as highlighted by Morgan et al [22].

There are other advantages of using open source competition such as cost, sustainability and quality as highlighted by Morgan et al [22]. The participants were provided with the original CAD model and initial stress analysis data in Fusion 360[20] to help in the initial re- design process.Data capture: The goal was to capture the visual behavioral activities using eye tracking and screen recording equipment. The FOVIO FX3 screen-based eye tracker from EyeTracking Inc was used to observe visual attention of participant using the gaze point (red colored square) in case., ... __,_ .. _, . .,,, .... .'-"-...... --"----"-'---' .,,_ .,_ ,111- .,_ ·- ,, ... -,_ .... , c. ... ·-•Qil0.0-1 ' __ ,......., ,.,_ ·•- ... - ,._ ·-c-" L-• ·-m-::::::--l l l ~ .......... .._C.n_,,CJ- - Ill_, • iilil c.,._, o- •f'.t(tQIIII<, .. -...-- •QtCIOOl.1« .. _ _ _ [!I ~ ... ,CJ- - Ill_, • iilil c.,._, o- •f'.t(tQIIII<, .. -...-- •QtCIOOl.1« .. _ _ _ [!I ~ ...

.._C.n_,,CJ- - Ill_, • iilil c.,._, o- •f'.t(tQIIII<, .. -...-- •QtCIOOl.1« .. _ _ _ [!I ~ ... ,CJ- - Ill_, • iilil c.,._, o- •f'.t(tQIIII<, .. -...-- •QtCIOOl.1« .. _ _ _ [!I ~ ... ,Cl=- 0 -"•• •o- •vo'"".l f ~ i► CJ--· ~ ,~.._--·~;,. : : m~~multiple windows are used on the same screen as shown in Figure 1. For the 90-minute re-design activity, participants were also instructed to indicate the build orientation of the part and were advised to avoid using lattice structures for light weighting.

For the 90-minute re-design activity, participants were also instructed to indicate the build orientation of the part and were advised to avoid using lattice structures for light weighting. The use of lattice structures would have diverted the attention of participants from using the shape complexity design freedom offeredFigure 1 : With the use of gaze tracking, the participants’ attention patterns were identifiedby AM to hierarchal complexity which was not desired [23].Analysis of Human Design Behaviors: Analysis of the screen capture data occurred through qualitative data analysis methods for real-time data, as developed and validated by the research team in past literature for other observational data in engineering education research contexts [24-25]. The corpus of data to be analyzed comprises six sets of logged CAD and visual activities representing the design processes of the six engineering student participants.

The corpus of data to be analyzed comprises six sets of logged CAD and visual activities representing the design processes of the six engineering student participants. Consistent with qualitative methodological traditions in engineering education literature and design thinking literature, behaviors can be sorted into representative functions such that each behavior could fit into a more generalizable theme, grouped with similar behaviors using well-established methods for the constant comparative method proposed originally by Glaser and Strauss [26] and well- accepted across all disciplines who employ any qualitative data analysis [27]. The first step in69qualitative categorization is to develop a “codebook” through constant comparative methods to define a comprehensive set of behaviors, which are also known as codes. Open and axial coding methods allow researchers to group codes into overarching themes.

Open and axial coding methods allow researchers to group codes into overarching themes. In our case, we used a combination of a priori and emergent coding methods to develop themes, employing standard language from the basic functional use of SolidWorks features combined with researcher descriptions of the participants’ attention patterns.Table 1 depicts the codebook for this data to describe the screen recorded and eye tracking behavioral data captured from the six participants. The different kinds of activities and spatial attention focus were grouped into three major categories of verification, composition, and modification. The verification category included the span of time which was spent by the participant on stress analysis, reading the design prompt and visual inspection of dimensions and geometry for AM feasibility. Activities like considering overhang angles, support structures and build orientation were included in the inspection category.

Activities like considering overhang angles, support structures and build orientation were included in the inspection category. The stress analysis category included visual attention of participants when they are observing initial FEA results provided in Autodesk Fusion 360 in addition to carrying out iterative FEA analysis on the geometry re-designed by them. The composition group includes the time spent by participants using extrude and sketch features in SolidWorks primarily used for creating a new geometric feature. Use of features like smoothing, fillet and revolving were categorized in the revising group under modification category. Making changes to the existing sketch or new sketch created by the participant was also included in the revising group. Editing of existing and newly created sketches and changing/scaling dimensions of geometries was included in the editing sub-category.

Activities which include eliminating and removing material using cut/extrude features are included in the remove material group.Table 1: Codebook for qualitative data analysis methodsOVERARCHING THEMEBEHAVIORAL ACTIVITY “CODE”DESCRIPTIONCODE NUMBERStress AnalysisObserving FEA results and performing stress analysis on created component1VerificationRequirement/Design promptFocusing on problem and objectives2CompositionModificationInspectionAdd material SketchingRevisingEditing Remove materialInspecting dimensions and considering AM restrictions Adding new features Sketching Smoothing existing features and using Fillet function Editing sketch, changing dimensions Removing existing features34 567 8For the purposes of developing time stamp and frequency data of each activity, we then assigned each code a numerical value for ease of data processing in MATLAB and MS Excel.

The numerical values have no significance on importance or order (e.g., category 1 is not superior or inferior to category 3), but are useful for computational bookkeeping purposes, a method applied in other studies [28-30]. The behavioral data, represented by numerical values, was then be70analyzed as a function of the percentage of the total time spent which is further discussed in the results section.Quantifying Design Manufacturability. After each participant completed the design challenge, the study aimed to quantify the design quality of each of the re-designed CAD model based on the primary criteria of weight, build time estimate, total volume, support volume and strength expressed via factor of safety. The support structures are generated using the standard SLM parameter set with a support critical angle of 35° on Autodesk Netfabb [31]. The parts were repaired using the extended repair script from Netfabb and then loaded into the EOS M290 machine workspace.

The parts were repaired using the extended repair script from Netfabb and then loaded into the EOS M290 machine workspace. The build strategy of EOS Print Standard Parameters set 30 microns for Stainless Steel 316 available in Netfabb was selected. The parts were raised by 1.5mm above the build platform to account for the wire EDM process. Once the build was ready with the support structures, the slice data was exported to ATLAS 3D [32] to simulate the laser powder bed fusion build process to predict thermal distortion and possibility of re-coater interference. The EOS M290 SS316 L parameters were selected for simulation on ATLAS 3D. The entire process workflow was repeated for all six participants as depicted in Figure 2. The manufacturability matrix table is generated using part details, build details and simulation metrics of all six participants. This matrix is specific to the laser-powder bed fusion process onlyExport CAD file in.

This matrix is specific to the laser-powder bed fusion process onlyExport CAD file in . step using SolidworksImport CAD file in Autodesk i. Netfabb Ultimate and perform( Start } ( End )-- and recoater inteference - indicated byrepair action using extended scriptRecord thermal distortionSelect orientation asdataparticipant... M290 machineLoad part into EOSworkspace... support files intoImport part andATLAS3Df-,11structures using SLM -Generate supportscriptExport part andsupport seperately into I.,_st!Figure 2 : Data analysis workflow for participant-generated CAD fileLimitations: As with any study, there are limitations due to study constraints. In particular, we acknowledge that this is a very small sample size; however, even six participants yielded several hours of data to analyze. Preliminary analysis of this data is also required in order to accurately pursue and analyze larger data sets with more participants.

Preliminary analysis of this data is also required in order to accurately pursue and analyze larger data sets with more participants. Another limitation is the population: While these are experienced designers who have DfAM formal education, they may not be experts in these areas. The re-design activity required a certain level of proficiency in CAD, which is not the same for all participants and may have affected performance in the re-design challenge. The categories in which the behavioral activities are divided is not based on any proven model and hence further validation is required.

The categories in which the behavioral activities are divided is not based on any proven model and hence further validation is required. Lastly, the analysis and coding for this paper was accomplished by a single researcher, such that in more robust studies, interrater reliability will need to be calculated to as one way of establishing quality in qualitative data.ResultsThe findings for this study are discussed first by analyzing the quality of the re-designed AM part using the criteria for manufacturability as listed in the Methods section. The manufacturability index will then be used to compare the designs across participants. Armed with71this information, the designers’ individual behaviors will be analyzed with respect to the various performance of their designs in terms of manufacturability.Manufacturability Matrix for Participants’ Final Designs.

The criteria for manufacturability (Weight, build time, volume, support volume, recoater interference, thermal distortion, and strength factor) were compared for all participants and compared with the original traditionally manufactured part that the participants were challenged to re-design.

All redesigns from the six participants resulted in a decreased part volume, and most resulted in a decreased support volume. The other criteria can be compared by inspection.Comparison of manufacturability index for participants. Since there is no specific index or ranking system established in academic literature to quantify the manufacturability of parts for AM, we adopted a normalization approach based on Marler et al. [12], where metrics of from all other participants are compared with the best performing participant in each category and the derived value is therefore normalized. The normalized stacked bar chart is shown in Figure 3.

The normalized stacked bar chart is shown in Figure 3. The longer the bar of a participant in a certain category, the better the performance of the design in that manufacturability criteria.72Normalized Manufacturability IndexBuild time estimate■Volume■Support Volume■SW Ratio Weight■■60.620.440.770.970.44r e b m u N540.730.550.360.560.710.960.750.910.370.55t n a p i c i t r a P320.711.000.540.651.000.840.970.551.001.0010.990.931.000.930.93Figure 3: Normalized manufacturability analysis comparing participant designsAs shown in Figure 3, the behavioral activities of each participant are ranked in comparison with the best in that category ranked as 1.00. For example, Participant 2 has the lowest build time estimate compared to all other participants and a lower build time results in a more favorable design for AM.

For example, Participant 2 has the lowest build time estimate compared to all other participants and a lower build time results in a more favorable design for AM. The design generated by Participant 2 has the lowest build time estimate, least overall and support volume, highest strength to weight ratio and lowest weight reduction; therefore, distinctly performing best in this design challenge. In contrast, the design created by Participant 4 has the shortest bar length for all performance categories (except support volume) which renders it to be the least favorable design for manufacturability.Aggregate Analysis of Designer Behaviors during Design Challenge. The quality of a particular build is interesting with respect to the proportion of total time each designer spent on a given code (i.e., a given behavior or category as per our qualitative codebook).

The quality of a particular build is interesting with respect to the proportion of total time each designer spent on a given code (i.e., a given behavior or category as per our qualitative codebook). To visualize the aggregate view of the proportion of total time spent by each designer on a particular activity, we employed content analysis methods to quantify the qualitative data collected from participants in this study. As an example, if a designer’s process involved spending 10 minutes on stress analysis out of a total of 100 minutes, the behavior would be plotted at the 0.1 mark for that participant, thereby normalizing the plots of all six designers’ design processes. Figure 4 shows a line plot of the codes from each of the participants, showing that stress analysis and sketching are the two major categories where participants spend their time.

Future work with a larger sample size of participants will yield statistical conclusions, and may point to indications that combinations of behaviors, or a certain pattern of occurrences within the design task that may matter to manufacturability.DiscussionTo the best of our knowledge, this is one of the first studies in the design (or re-design) for AM literature that discusses the role of designer behavior on the manufacturability and efficacy of the final additive design. As shown in our results, the participants who ranked the highest in manufacturability exhibited some of the same characteristics, namely, significant attention on stress analyses rather than other behaviors. In contrast, the participants who generated low performing designs spent a great deal more time on sketching rather than stress analysis.Although this is a small sample size, there are implications from this research that will inform future research directions and practice in the DfAM body of knowledge.

First, AM education should focus on developing designers’ habits of mind to focus first on the activities that result in a higher performance. Based on our preliminary results, that would mean reminding students to spend more time on stress analysis than sketching or other more intuitive design tasks.Based on the findings from this study, we suggest a designer-centric workflow to teach effective re-design processes for AM, focusing on designer behaviors. A relatively simple approach of re-designing a part, specifically for the laser powder bed fusion process is proposed as shown in Figure 5 based on this empirical study.

This workflow is valid for re-designing a single part with fewer number of loading and boundary conditions where designers engineering intuition can be used for deciding the optimal material distribution.Loading conditionsStandard lloy/material propertiesOptimization criteria/Max stiffnessBuild Orientation!Import file in CAD package~FEA analysisMaterial removal using primitivesRe-designed CAD fileYes.-Target Factor of safetyDefine design spaceNFigure 5: Proposed workflow to encourage behaviors that correspond with manufacturability in Re-design for Additive Manufacturing processesIn future work, we plan to validate the effectiveness of this process flow-chart in by extending our study to a larger number of participants in order to understand the statistical effects of spending more time on one activity than another.

The application of this process workflow will require preliminary understanding of opportunistic and restrictive design considerations of the laser powder bed fusion process. The re-design process chart can also be used by Additive manufacturing design educators for an introductory exposure to the re-design process. Other future work includes the advanced analysis of time-resolved design data from participants to elicit valuable heuristics that can optimize designer behaviors and education in industry and academic75settings; evaluate the effect of behaviors, combinations of behaviors, and occurrence patterns with manufacturability; and conducying comparison studies between expert AM designers with novices.

In this way we intend to systematize and characterize the art of DfAM to translate more effectively across sectors interested in Addive Manufacturing.ConclusionThis empirical study investigated the design processes of six graduate-level engineering designers specializing in additive manufacturing as they were challenged to re-design a traditionally manufactured part to be optimal for Additive Manufacturing. The participants’ decisions were captured using screen capture and eye tracking methods, as well as the action log of design software. The behaviors of the designers were qualitatively coded and compared with the final design efficacy, measured in terms of manufacturability based on several key parameters. Our findings indicate that the participants who designed the most effective designs spend more time performing stress analyses on their designs.

Our findings indicate that the participants who designed the most effective designs spend more time performing stress analyses on their designs. Implications from this study, if upheld by future work with a larger sample size, indicate that DfAM education might benefit by guiding designers to focus on activities that have a more substantial impact on design quality. To our knowledge, this is one of the first studies that links AM designers’ behaviors to manufacturability in the context of design and re-design for additive manufacturing.Design challengeAppendix A: Design ChallengeAppendix B: Example CAD ScreenshotsObjectiveThe objective of this challenge is to redesign the ALCOA bearing bracket in such a way that its topology and shape are optimized for minimizing weight while fitting in the target envelope and meeting the technical requirements.

The bracket is intended to be additively manufactured (using laser powder bed fusion technique) and the design shall also minimize and/or eliminate the need for support structures. The efficacy of the design submission will be eva luated via FEA, strength-to-weight ratio and manufacturability.Original designAlcoa Bracketa) You are provided with a CAD design of the existing model b) Based on intuition and results from FEA analysis, you can come up with a design with minimumc)weight/volume to meet the given loading conditions.

It is advised to concentrate on bu lk removal of materia l and not use lattice structures for light weighting the part design.d) Please indicate the intended print orientation/direction in your submission along with one-twosentences of justification for your choice.e) The maximum time duration for this activity is 90 minutesDesign require mentsDesign material: 15-SPH per AMS5862: Elastic Modu lus (E) = 29,000 KSI = 200,000 MPa = 200 GPa Poisson Ratio (v) = 0.27 Yield Stress (o_y) = 145 KSI = 1000 MPA Density (p) = 0.283 lb/in'3 = 7833 kg/m'3 Material is assumed to be linear elastic Minimu m geometric feature: 0.025 in.

Minimum wa ll thickness: 0.045 in.Participant 2 re-Parts shall be optimized for minimum weight with the following boundary and loading conditions: - Base support: The part is bolted against a mating plate of high stiffness - Bolts interface: The parts is fastened with four #10-32 high strength tension rated bolts as indicated in the specifications - Bearing interface: The part is loaded through a high stiffness spherical bearing with three load cases: 1. A load of 1,250 lbf applied horizontally 2. A load of 1,875 lbf applied 45 degrees from the horizontal 3. A load of 2,500 lbf applied vertically76References[1]S. Yang and Y. F. Zhao, “Additive manufacturing-enabled design theory and methodology: a critical review,” Int. J. Adv. Manuf. Technol., vol. 80, no. 1–4, pp. 327– 342, 2015.[2] W. Yan et al., “An integrated process–structure–property modeling framework foradditive manufacturing,” Comput. Methods Appl. Mech. Eng., vol. 339, pp. 184–204, 2018.[3]Z.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEnabling Cost-based Support Structure Optimization in Laser Powder Bed Fusion of MetalsK. Bartsch1 and C. Emmelmann11Institute of Laser and System Technologies, Hamburg University of Technology, Hamburg, GermanyAbstractSupport structures are essential to laser powder bed fusion (PBF-LB/M). They sustain overhangs, prevent distortion, and dissipate process-induced heat. Their removal after manufacturing is required, though, increasing the overall costs. Therefore, optimization is important to increase the economic efficiency of PBF-LB/M. To enable optimization focused on the support structures’ costs, a cost model is developed. The whole production process, including the design, manufacturing, and post-processing of a part is considered by deriving formulas for the individual costs.

The whole production process, including the design, manufacturing, and post-processing of a part is considered by deriving formulas for the individual costs. The cost model is applied to a benchmark procedure previously developed. Additionally, a case study investigating different support layout strategies is conducted.The full paper may be found in a special issue of the TMS publication JOM, March 2022.240.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperENERGY AND EMISSION ESTIMATION UNCERTAINTY IN FUSED DEPOSITION MODELING FOR A JOB-SHOPLee M. Clemon*†Department of Mechanical Engineering, University of California, Berkeley, CA 94709 †Sandia National Laboratories, Livermore, CA 94551Sandia National Laboratories is a multi-program laboratory managed and operated by Sandia Corporation, a wholly owned subsidiary of Lockheed Martin Corporation, for the U.S. Department of Energy's National Nuclear Security Administration under contract DE-AC04-94AL85000.AbstractSolid freeform fabrication has the potential to affect both financial and environmental concerns for manufacturing enterprises.

However, when planning for installation of a new machine tool, accurate energy usage estimation relies heavily on the data and model selections of the estimator. This project used a variety data sources and model decision options to examine the spread of energy consumption and global warming potential estimates for a fused deposition modeling machine. In addition to primary and secondary data sources, the use of similar machines was explored as proxy estimates for the target machine. A Monte Carlo simulation was constructed to vary the model selections, machine utilization, and data sources. The results indicated data sources and model decisions had large effects on the output and that most model estimates were low.IntroductionThe manufacturing industry is ever-growing more conscious of the triple-bottom line (financial, environmental, and social impacts) of production activities.

In addition to awareness, increased instrumentation of machine tools provides data to better assess each impact area. One commonly employed technique to evaluate any, or all, of these impact areas is Life Cycle Assessment (LCA), where the inputs and outputs of each process in each life cycle phase are tabulated [1]. The tabulated results typically provide point estimate of the impact, but often lack sufficient information about uncertainty in the result to be widely useful. Although many methods have been proposed to address uncertainty issues in LCA point-estimates, no clear solution has been accepted in the literature [2] [3] [4] [5].Uncertainty in this document is categorized into three areas 1) parameter, 2) model, and 3)scenario, following [6]. First, parameter uncertainty arises from inaccuracy of the value(s) assigned to each variable, or parameter, in a given model.

First, parameter uncertainty arises from inaccuracy of the value(s) assigned to each variable, or parameter, in a given model. Parameter uncertainty stems from inherent randomness of the process, assignment of data from proxy technologies/machines, error in collected test data, and non-representative data for the population in question. Second, model uncertainty is attributed to simplification of process physics, negligence of pertinent processes or inputs, and incomplete or inaccurate process knowledge. Third, scenario uncertainty comes from the use of aggregate data which assumes homogeneity and sourcing data from inadequately representative processes or exclusion of relevant data. Scenario uncertainty could be ignoring pertinent the geographical, temporal, or technological specifics of the system in question.1878When uncertainty is included in a LCA, typically only parameter or scenario uncertainty have been considered.

Some proposed qualitative methods for evaluating input data sources [7] [8], but fail to add significant value in comparing results or quantitative risk assessment. More robust methods for incorporating uncertainty involve statistical simulation [9] [10] [11], but typically only vary the values of predetermined parameters in a single fixed model. Many LCA studies do not address the issue of uncertainty, some avoid interpretation errors [12]. Thus, a more comprehensive inclusion of uncertainty in LCA is needed to ensure usefulness of the results. Namely, a statistical simulation that includes model variations can illustrate the variety of estimates that may be generated for a single assessment. Environmental impacts are often reported using a variety of metrics which represent the potential for environmental damage.

Environmental impacts are often reported using a variety of metrics which represent the potential for environmental damage. In this work, only gaseous emissions that contribute to climate change were considered, historically tabulated as carbon-dioxide equivalent mass and named Global Warming Potential (GWP) [13]. The recent interest in Additive Manufacturing and 3D printing, but limited work assessing sustainability, makes the case study on Fused Deposition Modeling (FDM) in this paper highly relevant.The objective of this study was to examine the combined effects of parameter, scenario, and model uncertainty in predicting the energy usage and associated global warming potential of parts made on a FDM machine in a Northern California job shop.MethodologyThis methodology is separated into two main sections the first describes the prediction scheme and second describes data collection.

The prediction scheme includes how the simulation of energy intensity and GWP estimates was conducted, the models. The prediction scheme allowed for data at different levels of granularity, scenarios. For example, the U.S.A. national average cost per kilowatt-hour of electricity was less precise of a proxy than the California average for estimating energy costs for a machine installed in Berkeley, California. Since the actual implementation details of a target machine are not necessarily known in a general application, proxy estimators were included. The data collection section provides how power measurements were taken and what literature data was included.Prediction SchemeA variety of scenarios may exist for the installation and operation of a given machine tool.Similarly, literature estimates for a single machine tool implementation may encode these differences due to investigator knowledge, literature completeness, devoted study time, and invested resources.

To model this variety, several combinations of data sources represented varying degrees of 1) scope of included elements, 2) proxy appropriateness, and 3) investment in the estimate. The scope of an estimate determines which major components are included, here major elements were denoted Value Add Process, Machine Auxiliary Processes, HVAC, and Lighting. Each iteration of the scope in the simulation was expanded by including an additional component (Table 1). The selected proxies were chosen to represent plausible data sources of similar technology, another FDM machine, and a photopolymer-jetting machine. Each unique combination of the data sources represents a single scenario. Practitioners may favor different models and data sources, this was captured by iterating over several options within each defined scope (Table 1). To capture this variety, at each Scope, common models were considered1879separately for each Component (Table 2).

To capture this variety, at each Scope, common models were considered1879separately for each Component (Table 2). Each unique combination of Scope, Scenario Model, and Data Source represents a set of point estimates.While the Scenario Models in Table 2 provide a high level description of the calculation,the details follow. Energy was defined as the integral of instantaneous power over time. This was discretized into a sum over the various contributing energy sources, listed as Components in Table 1 and Table 2. For generality, the energy delivery efficiency, denoted into extraction from raw materials, transportation of the extracted materials, conversion into electricity, and finally transmission to the machine. However, since all of the machines considered in this study received energy from the same system this delivery efficiency was extracted from the sum as a constant, to the sum of instantaneous power consumption, (Equation (1)).

This was further broken down to include the energy consumed by the machine,, was broken down𝜂𝜂. Energy per kilogram of part produced is equivalent, over the time,, required to produce the part𝜂𝜂total𝑃𝑃 (Equation (1))., the heating, ventilation and air-conditioning (HVAC),allocated to the machine, 𝐸𝐸𝑚𝑚𝑚𝑚𝑚𝑚ℎ𝑖𝑖𝑖𝑖𝑖𝑖𝑡𝑡𝐸𝐸𝐻𝐻𝐻𝐻𝐻𝐻𝐻𝐻, and the lightingElighting(1)𝐸𝐸 ∶= ∫ 𝑃𝑃𝑃𝑃𝑡𝑡 ≅ ∑𝑃𝑃𝑖𝑖𝑡𝑡𝑖𝑖 𝜂𝜂𝑖𝑖 =�𝐸𝐸𝐻𝐻𝐻𝐻𝐻𝐻𝐻𝐻+𝐸𝐸𝑚𝑚𝑚𝑚𝑚𝑚ℎ𝑖𝑖𝑖𝑖𝑖𝑖+𝐸𝐸𝑙𝑙𝑖𝑖𝑙𝑙ℎ𝑡𝑡𝑖𝑖𝑖𝑖𝑙𝑙� �𝜂𝜂𝑡𝑡𝑡𝑡𝑚𝑚𝑖𝑖𝑡𝑡𝑚𝑚𝑖𝑖𝑡𝑡𝜂𝜂𝑚𝑚𝑐𝑐𝑖𝑖𝑐𝑐𝜂𝜂𝑡𝑡𝑡𝑡𝑚𝑚𝑖𝑖𝑡𝑡𝑡𝑡𝑐𝑐𝑡𝑡𝑡𝑡𝜂𝜂𝑖𝑖𝑒𝑒𝑡𝑡𝑡𝑡𝑚𝑚𝑚𝑚𝑡𝑡�The Value Add process is denoted production in the subscripts of the equations(Equations (2) and (3)). This production and additional machine processes for FDM and Photopolymer Jetting were discretized into a warm-up, production, idle, and standby operational phases with differing time lengths following the work of [14]. Each operational phase duration was scaled by the total of the phases to provide a utilization, or duty cycle, for the machine.

Each operational phase duration was scaled by the total of the phases to provide a utilization, or duty cycle, for the machine. The time to complete a given job is also varied in accordance with the range of job sizes collected over the course of the sample period, denoted, with the average length. This duration ofthen provided the length of time needed to deposit one kilogram of material based on the tjob previously determined utilization. Adjusting the time lengths in this manner was to apply the tjob energy used during down time of the machine to the products it produced.𝑡𝑡̅𝑗𝑗𝑗𝑗𝑗𝑗(2)𝐸𝐸𝑚𝑚𝑚𝑚𝑚𝑚ℎ𝑖𝑖𝑖𝑖𝑖𝑖 = �𝑃𝑃𝑡𝑡𝑡𝑡𝑐𝑐𝑝𝑝𝑝𝑝𝑚𝑚𝑡𝑡𝑖𝑖𝑐𝑐𝑖𝑖𝑡𝑡𝑡𝑡𝑡𝑡𝑐𝑐𝑝𝑝𝑝𝑝𝑚𝑚𝑡𝑡𝑖𝑖𝑐𝑐𝑖𝑖+𝑃𝑃𝑖𝑖𝑝𝑝𝑙𝑙𝑖𝑖𝑡𝑡𝑖𝑖𝑝𝑝𝑙𝑙𝑖𝑖+𝑃𝑃𝑡𝑡𝑡𝑡𝑚𝑚𝑖𝑖𝑝𝑝𝑠𝑠𝑠𝑠𝑡𝑡𝑡𝑡𝑡𝑡𝑚𝑚𝑖𝑖𝑝𝑝𝑠𝑠𝑠𝑠+𝑃𝑃𝑤𝑤𝑚𝑚𝑡𝑡𝑚𝑚−𝑝𝑝𝑡𝑡𝑡𝑡𝑤𝑤𝑚𝑚𝑡𝑡𝑚𝑚−𝑝𝑝𝑡𝑡𝑡𝑡𝑡𝑡𝑐𝑐𝑡𝑡𝑚𝑚𝑙𝑙𝑡𝑡𝑗𝑗𝑗𝑗𝑗𝑗�1 𝜂𝜂𝑡𝑡𝑐𝑐𝑡𝑡𝑚𝑚𝑙𝑙The ambient temperature control (HVAC) and lighting in the facility were allocated to all of the machines according to the models listed in Table 2 with data from [15] [16] (Equation (3)).

Where the Square Footage model means the total power used by the metered HVAC system was applied equally over the area used by the machine and auxiliary components. The Thermal Load model used the waste heat, assumes the HVAC system must counter this energy load (Equation (3))., generated by the machine during each operational phase and𝑄𝑄𝐻𝐻𝑚𝑚𝑚𝑚𝑚𝑚ℎ𝑖𝑖𝑖𝑖𝑖𝑖+𝐻𝐻𝑚𝑚𝑝𝑝𝑒𝑒𝑖𝑖𝑙𝑙𝑖𝑖𝑚𝑚𝑡𝑡𝑖𝑖𝑖𝑖𝑡𝑡 𝐻𝐻𝑚𝑚𝑖𝑖𝑡𝑡𝑖𝑖𝑡𝑡𝑖𝑖𝑝𝑝𝑃𝑃𝐻𝐻𝐻𝐻𝐻𝐻𝐻𝐻𝑡𝑡̅𝑗𝑗𝑗𝑗𝑗𝑗1 𝜂𝜂𝑡𝑡𝑐𝑐𝑡𝑡𝑚𝑚𝑙𝑙 �𝐸𝐸𝐻𝐻𝐻𝐻𝐻𝐻𝐻𝐻 =𝑄𝑄𝑡𝑡𝑡𝑡𝑐𝑐𝑝𝑝𝑝𝑝𝑚𝑚𝑡𝑡𝑖𝑖𝑐𝑐𝑖𝑖𝑡𝑡𝑡𝑡𝑡𝑡𝑐𝑐𝑝𝑝𝑝𝑝𝑚𝑚𝑡𝑡𝑖𝑖𝑐𝑐𝑖𝑖+𝑄𝑄𝑖𝑖𝑝𝑝𝑙𝑙𝑖𝑖𝑡𝑡𝑖𝑖𝑝𝑝𝑙𝑙𝑖𝑖+𝑄𝑄𝑡𝑡𝑡𝑡𝑚𝑚𝑖𝑖𝑝𝑝𝑠𝑠𝑠𝑠𝑡𝑡𝑡𝑡𝑡𝑡𝑚𝑚𝑖𝑖𝑝𝑝𝑠𝑠𝑠𝑠+𝑄𝑄𝑤𝑤𝑚𝑚𝑡𝑡𝑚𝑚−𝑝𝑝𝑡𝑡𝑡𝑡𝑤𝑤𝑚𝑚𝑡𝑡𝑚𝑚−𝑝𝑝𝑡𝑡𝑡𝑡̅𝑗𝑗𝑗𝑗𝑗𝑗� Once the scenarios were defined, a Monte-Carlo simulation was then conducted using�𝑡𝑡𝑡𝑡𝑐𝑐𝑡𝑡𝑚𝑚𝑙𝑙(3)MATLAB. For each scenario, each parameter from each data set was conditioned with a1880probability distribution. Since many of the available data sources provided either a mean and standard deviation, or just the range of values the probability distribution is not known.

Since many of the available data sources provided either a mean and standard deviation, or just the range of values the probability distribution is not known. Thus, probability distributions were selected among Uniform, Gamma, Normal, and Lognormal as an additional model variation (Table 3). When mean and variance are provided, the maximum and minimum values were set to 5 standard deviations away from the mean value. Similarly, when only a maximum and minimum value were provided, the mean was set as the average of the two values and the minimum set as 3 standard deviations below the mean. Fixing the relations of minimum value and mean value allowed for calculation of probability distribution factors where appropriate.Table 1. Resource consumption analysis boundary conditions, or scopes.

The parameters were then combined according to the selected model for each component and the energy and GWP values totaled. GWP was calculated by multiplying the energy values by the weighted sum of GWP factors for the selected data set. Different energy sources had different GWP factors and they were combined prior to multiplication. The majority of the simulation used a single probability distribution for all of the parameters, selected from Table 3. One additional distribution of estimates was generated making selections more specific to the ‘target’ machine. This target machine served as a datum to compare all other estimates and represents the narrowest bounds on the estimated results.Data CollectionPower Consumption MeasurementEnergy consumption was efficiently collected by characterizing the machine tool energyusage by operation phase. A Yokogawa CW240 placed in between the machine power cord and the wall outlet on the shop floor.

A Yokogawa CW240 placed in between the machine power cord and the wall outlet on the shop floor. Each machine was used normally in the course of the machine shop business. During the course of operation output mass and time of production were recorded for 6 weeks. The process for this energy characterization was inspired by the Baseline Energy Consumption model from [17]. Clemon, et al., adapted this model for additive manufacturing in FDM and Photopolymer-jetting machines [14] [18]. Thus, the amount of data required to determine the energy consumption of each machine was greatly reduced without significant loss of granularity. Machine operation was characterized into several operational phases. The operational phases for sorting energy data include warm-up, idle, production, and standby. For each build, energy demands were averaged based on operation phase. The variation recorded in the data provided a variance for demand in each phase.

For each build, energy demands were averaged based on operation phase. The variation recorded in the data provided a variance for demand in each phase. Further, the minimum and maximum uptime for the machine over the recording period provide a percentage and range of utilization for the simulation. Uptime is collected from the machine on-board job log.Environmental Impact EstimateThree environmental impact scenarios were considered for Global Warming Potential (GWP) emissions, measured by carbon dioxide equivalent weight for greenhouse gases, due to energy consumption. The first emissions scenario, denoted G1, was the estimated impact according to an Economic Input-Output (EIO) analysis. EIO analysis uses information from the U.S. Commerce and U.S. Environmental Protection Agency in concert for a linear estimate of the GWP per kWh purchased from the power sector [19].

EIO analysis uses information from the U.S. Commerce and U.S. Environmental Protection Agency in concert for a linear estimate of the GWP per kWh purchased from the power sector [19]. The second GWP scenario, G2, used data on the upstream and downstream externalized effects of power delivery from a report by the National Academy of Sciences [20]. This report provided energy delivery efficiency for both G1 and G2. In G2, estimated GWP emissions and energy by fuel type were retrieved from [20], the mix of fuel types was determined by the selected energy use data for each scenario. Finally, the third GWP scenario, G3, used a combination of literature for the California specific energy mix and more thorough accounting of GWP emissions from [20] [22] [21].Error AssessmentIn order to assess the error in a simulated estimate, details of the final machine tool installation were required.

The machines studied were Dimension 1200SST produced by Stratasys for FDM, uPrint SE produced by Stratasys for FDM, and Connex350 produced by Objet for Photopolymer-Jetting. The Dimension 1200SST was selected as the target machine, or datum, for subsequent comparison and uncertainty calculation.1882Two proxy machines were used to simulate the estimation process, uPrint SE andConnex350. The uPrint SE was expected to be a more accurate proxy given it was also an FDM machine, the Connex350 serves as a proxy via an adjacent technology. Operational scenarios were selected based on 6 weeks of monitoring the frequency and duration of use of all 3 machines in a university job shop in Berkeley, California. Data on the facility HVAC and lighting energy consumption was collected from utility monitoring and added as an additional input to energy consumption.

Data on the facility HVAC and lighting energy consumption was collected from utility monitoring and added as an additional input to energy consumption. Peripheral components used to run the machines, such as computers, were considered captured in the building usage since these machines support a large number of activities in addition to the target machine.The environmental impact scenario options are limited to GWP as measured in kilogramsof carbon-dioxide equivalent weight for this study. Data sources for energy mix, delivery efficiency, and GWP produced per kilowatt-hour began most broadly with economically correlated emissions and improve in specificity with literature data from national and state- specific sources. EIO data serves for the most basic and broad estimate scenario [19]. The intermediately detailed estimate data comes from the National Research Council [20].

EIO data serves for the most basic and broad estimate scenario [19]. The intermediately detailed estimate data comes from the National Research Council [20]. The most detailed accounting of emissions and efficiency comes from a combination of literature, namely the State of California [21], the National Academy of Sciences [20] and a quality study by Hondo [22].ResultsThe inclusion of various data sources and energy allocation methods into the predictionscheme provided an informative distribution of energy and global warming potential estimates. Distinct operational phases were present for all three machines. Separation of operational phases was completed by manually selecting transition points in accordance with the machine controller display and checked graphically.Table 4.

Separation of operational phases was completed by manually selecting transition points in accordance with the machine controller display and checked graphically.Table 4. Power demand by operation state for Dimension 1200SST Fused Deposition Modeling Machine Average Power (W) 685 243 784 162Standard Deviation (W) 270* 63 26 23Operation StateWarm Up Idle Production StandbyPower demand for the Dimension 1200SST Fused Deposition Modeling machine clearlyshows differing low frequency cycles during different operating phases. The warm up phase has a long period of continuous power demand preceded by a short energy spike and short period of low power demand. The production phase has a low cycle of high energy demand as well as high frequency oscillations with lower amplitude (Figure 1).The anomaly was induced by the loading door being left open for an extended period of time (5 minutes). A short secondary warm-up period is contained within the anomaly.

A short secondary warm-up period is contained within the anomaly. Summary data for each phase indicates the productionDistribution for warm-up is multi-modal contributing to a large variance1883phase has the highest overall power demand (Figure 1). Idle and standby phases are separated due to 1) idle may transition to production, but standby must transition to warm-up and 2) the differing power demand.Figure 1. Dimension 1200SST instantaneous power demand from off state with overnight soak and production of a small partFigure 2. uPrint SE instantaneous power demand from off state with overnight soak and production of a small part [14].1884The uPrint SE power demand graph follow very similarly to the Dimension 1200SST with a long warm-up period of continuous power consumption followed by a production phase with a distinct low frequency of high amplitude and higher frequency with lower amplitude Figure 2).

Differently than the two FDM machines, the Connex 350 has a constant high power consumption level when in production (Figure 3).Figure 3. Connex 350 instantaneous power demand from off state with overnight soak and production of a small part [14].The point estimates generated according to all of the variations in model, data source, andparameter values show a mean that is less than the specific case of the target machine (Figure 4). The mean energy consumption of the target machine was 92 kWh/kg with a standard deviation of 11 kWh/kg. The mean energy consumption of all estimates was 64 kWh/kg with a standard deviation of 15 kWh/kg. Three data sources were used to calculate GWP corresponding to the distribution of energy consumption estimates giving a tri-modal histogram. The majority of GWP estimates were below the target machine mean GWP estimate (Figure 5).

The majority of GWP estimates were below the target machine mean GWP estimate (Figure 5). The mean GWP of all estimates was 84 kg CO2-eq/kg part produced, with a standard deviation of 48 kg CO2- eq/kg part produced. The mean GWP estimate for the target machine was , 120 kg CO2-eq/kg part produced, with a standard deviation of 14 kg CO2-eq/kg part produced. The three peaks from all estimates result from the differences in each data set for GWP factors, whereas the target machine has a limited GWP factor range due to its location in Berkeley rather than anywhere in California, or the United States.1885Figure 4. Energy intensity estimates from all considered scenarios and data sourcesFigure 5.

Energy intensity estimates from all considered scenarios and data sourcesFigure 5. Global Warming Potential estimates from all considered scenarios and data sourcesDiscussionThe distribution of estimates indicates 1) not every part was produced with an equivalentembodied energy and 2) resource consumption and environmental impact estimates with minimal or missing variance information may be significantly different than the realized quantities.As mass customization becomes more commonplace with the rise of additive manufacturing, the methodology and results from this study suggest the varying embodied energy of every part may be uniquely identified. In particular, the connection of the duration scaling in Equation (2) to a single part could provide this insight. Interpreting a point estimate of energy usage or global warming potential may have significant and unrealized uncertainty based on the assumptions and decisions of the estimator.

Interpreting a point estimate of energy usage or global warming potential may have significant and unrealized uncertainty based on the assumptions and decisions of the estimator. This uncertainty is compounded when a point1886estimate for energy is then used to generate a point estimate for GWP as seen by the mass of estimates below the target machine in Figure 4 and again in Figure 5.The proposed method included all three types of uncertainty in a LCA in estimating theenvironmental impact due to energy consumption of a Fused Deposition Modeling machine in a Northern California job-shop. Implicit in the simulation construction was that every set of estimates is given the same weight from an equal number of samples, this suggests that a person chosen at random to complete an estimate is equally likely to construct any of the simulated results.

The results suggest that point estimates will tend to underestimate the impacts, and that proxy machines and processes are not necessarily good estimators for energy usage of a new technology. Indication of the potential error due to the use of different data sets–even from very similar machines–is demonstrated, which illustrates the need for this methodology. For example, two of the three datasets used for estimating GWP have a limited range and shift a large portion of the simulated estimates below the target. However, using the most liberal GWP factors from the U.S. national averages, the estimates shift above the target.The results support the notion that increased collection and clarity of data will continue to improve resource intensity predictions and environmental impact assessments. The use of data from proxy machines or technologies in an estimate may result in either overestimate or underestimate.

The use of data from proxy machines or technologies in an estimate may result in either overestimate or underestimate. Whether the estimation is over or under is not necessarily known a priori and dev. The decisions of the researcher in setting up an estimate of resource consumption and account for sub-processes has a significant effect on the outcome of the results. In appropriately assigned distributions may artificially widen or narrow the variance of an estimate. Even though life cycle assessments are intended to enable standardized comparisons, comparative analyses using multiple methods on the same data set are infrequently realized and should be conducted for completeness.The methodology presented here may be more broadly applicable to measuring and estimating resource intensities and environmental impacts of other additive manufacturing machines.

The energy characterization method from [14] and [17] has been shown useful for multiple FDM machines as well as Photopolymer Jetting, and could be useful in other additive technologies. Extension of this work to include material, human health, or other factors is possible with data and models for those factors. The results of this study demonstrate point estimates should be considered with great caution by decision makers and enterprise planning for resource consumption could benefit from estimates with wider ranges.References[1] International Standard Organization, ISO 14040: Enmental Management-Life CycleAssessment Principles and Framework, 1997.[2] J. Reap, F. Roman, S. Duncan and B. Bras, "A survey of unresolved problems in life cycleassessment," International Journal of Life Cycle Assessment, vol. 13, no. 5, pp. 374-388, May 2008.[3] S. Lloyd and R.

Roman, S. Duncan and B. Bras, "A survey of unresolved problems in life cycleassessment," International Journal of Life Cycle Assessment, vol. 13, no. 5, pp. 374-388, May 2008.[3] S. Lloyd and R. Ries, "Characterising, propagating, and analyzing uncertainty in life-cycle anssessment: Asurvey of quantitative approaches," Journal of Industrial Ecology, vol. 11, no. 1, pp. 161-179, 2007.1887[4] T. McKone and K. Bogen, "Uncertainties in health-risk assessment: an integrated casestudy based on tetrachloroethylene in california groundwater," Regulatory Toxicology and Pharmacology, vol. 15, no. 1, pp. 86-103, 1992.[5] S. Ross, D. Evans, M. Webber and , "How LCA studies deal with uncertainty," International Journal of Life Cycle Assessment, vol. 7, no. 1, pp. 47-52, 2002. [6] U.S. Environmental Protection Agency, Exposure Factors Handbook: 2011 Edition. Number 3.

23, no. 12, pp. 1461-1467, 2008.[10] S. Lo, H. Ma and S. Lo, "Quantifying and reducing uncertainty in life cycle assessmentusing the baysian monte carlo method," Science of the Total Environment, vol. 340, no. 1- 3, pp. 23-33, March 2005.[11] G. Sonnemann, M. Schuhmacher and F. Castells, "Unceratinty assessment by a montecarlo simulation in a life cycle inventory of electricity produced by a waste incinerator," Journal of Cleaner Production, vol. 11, no. 3, pp. 279-292, 2003.[12] A. E. Bjorklund, "Survey of approaches to improve reliability in LCA," The InternationalJournal of Life Cycle Assessment, vol. 7, no. 2, pp. 64-72, 2002.[13] A. Azapagic, A. Emsley and L. Hamerton, Polymers, the Environment and SustainableDevelopment, John Wiley & Sons, 2003.[14] L. M. Clemon, A. Sudradjat, M. Jaquez, A. Krishna, M. Rammah and D.

Azapagic, A. Emsley and L. Hamerton, Polymers, the Environment and SustainableDevelopment, John Wiley & Sons, 2003.[14] L. M. Clemon, A. Sudradjat, M. Jaquez, A. Krishna, M. Rammah and D. Dornfeld,"Precision and energy usage for additive manufacturing," in Proceedings for 2013 ASME International Congress and Exhibition, 2013.[15] U.S. Energy Information Administration, "Annual Energy Review," 2011. [16] University of California, Berkeley, "Energy Dashboards," University of California Regents, [Online]. Available: http://sustainability.berkeley.edu/energy-dashboards. [Accessed 16 March 2013].[17] A. Vijayaraghavan and D. Dornfeld, "Automated energy monitoring of machine tools,"CIRP Annals Manufacturing Technology, vol. 59, p. 189–201, 2010.[18] L. M. Clemon, Analysis of Uncertainty in Estimating the Resource Intensities and Impacts of Solid Freeform Fabrication Processes, vol. Masters Thesis, Berkeley, California: University of California, 2013.[19] L. Lave, H.

ERGY AND RESOURCE EFFICIENCY OF SLS/SLM PROCESSESK. Kellens*, E. Yasa*, Renaldi*, W. Dewulf†, J.P. Kruth* and J.R. Duflou*K.U.Leuven – Department of Mechanical Engineering, Leuven, Belgium † Group T – Leuven Engineering College, K.U.Leuven Association, Leuven, BelgiumAbstractManufacturing processes, as used for discrete part manufacturing, are responsible for a substantial part of the environmental impact of products, but are still poorly documented in terms of their environmental footprint. The lack of thorough analysis of manufacturing processes has as consequence that optimization opportunities are often not recognized and that improved machine tool design in terms of ecological footprint has only been targeted for a few common processes.Additive manufacturing processes such as Selective Laser Sintering (SLS) and Selective Laser Melting (SLM) allow near-net shape manufacturing of complex work pieces.

Consequently, they inherently offer opportunities for minimum-waste and sustainable manufacturing. Nevertheless, powder production, energy consumption as well as powder losses are important and not always optimized environmental impact drivers of SLS and SLM. This paper presents the results of a data collection effort, allowing to assess the overall environmental impact of these processes using the methodology of the CO2PE! (Cooperative Effort on Process Emissions in Manufacturing) initiative.Based on the collected LCI data, a subsequent impact assessment analysis allows indentifying the most important contributors to the environmental impact of SLS/SLM. Next to the electricity consumption, the consumption of inert gasses proves to be an important cause of environmental impact.

Next to the electricity consumption, the consumption of inert gasses proves to be an important cause of environmental impact. Finally, the paper sketches the improvement potential for SLS/SLM on machine tool as well as system level.KeywordsSelective Laser Sintering (SLS), Selective Laser Melting (SLM), Life Cycle Assessment, Sustainable Manufacturing, Energy and Resource Efficiency, CO2PE!-InitiativeIntroductionThe application of Solid Freeform Fabrication (SFF) techniques such as Selective Laser Sintering (SLS) and Selective Laser Melting (SLM) [1,2,3] has seen a remarkable expansion over the past ten years both in terms of parts produced as well as in the number of machine tools sold [4]. SFF techniques are often prized for being much cleaner than conventional machining processes and being able to fabricate objects with minimum waste.

SFF techniques are often prized for being much cleaner than conventional machining processes and being able to fabricate objects with minimum waste. Among others, further ecological advantages can be found in the reduction of transportation and supply chain impacts and the absence of special tooling or hazardous cutting fluids. However, quantitative analyses of the environmental impact caused by SFF processes are very limited and detailed data about environmental metrics such as energy consumption, waste flows and generated process emissions is often lacking [5,6]. Due to lack of thorough environmental analyses, optimization opportunities are often not recognized and improved machine tool design in terms of ecological footprint reduction has often not been targeted. In this paper, results of an environmental analysis of SLS as well as SLM production processes, using the CO2PE!-methodology [7,8], are presented. First a summarized description of theused methodology is provided.

First a summarized description of theused methodology is provided. Based on the results of the LCI data collection effort for SLS and SLM processes, the most important contributors to the environmental impact of SLS/SLM processes are identified by an impact assessment. Finally, the paper sketches the improvement potential for SLS/SLM on machine tool as well as system level.CO2PE!-methodologyIn order to enable a consistent and accurate inventorisation and documentation of Life Cycle Inventory (LCI) data, a generic data collection methodology [7], shown in Figure 1, is developed within the CO2PE!-Initiative [8]. While the functional unit and system boundaries are set during the goal and scope definition, the real data collection effort occurs based on industrial measurements and comprises four different studies. First a time study is performed in order to identify the different use modes of the processes.

First a time study is performed in order to identify the different use modes of the processes. Then power, consumables and emission studies are performed for all different modes on machine tool as well as subunit level.Figure 1: CO2PE!-methodology, adapted from [7]LCI Data CollectionThis section describes the Life Cycle Inventory (LCI) data collected using the above described methodology. The system boundaries, illustrated in Figure 2, of the analysis are set to include only the operating phase of the involved SLS/SLM equipment: EOSINST P760, EOSINT P360, EOSINT FORMIGA P100 [9] and Concept Laser M3 Linear machine [10].

First results of this data collection effort were already presented in [11].Figure 2: System boundaries (in- and outputs) Figure 2Selective Laser Sintering (SLS) Selective Laser Sintering (SLS)Time studyFor SLS, extensive time studies were performed on 3 different EOSINT P760 machine time studies were performed on 3 different EOSINT P760 machine time studies were performed on 3 different EOSINT P760 machine productive modes comprise machine tool cleaning, preheating (2 hours) tools [9]. The non-productive modes comprise machine tool cleaning, preheating productive modes comprise machine tool cleaning, preheating and cooling down (2 hours). The productive modes are responsible for ca. 87% of the . The productive modes are responsible for ca. 87% of the total . The productive modes are responsible for ca.

The productive modes are responsible for ca. 87% of the . The productive modes are responsible for ca. 87% of the total . The productive modes are responsible for ca. 87% of the production time, and can be subdivided further into three main modes: the laser exposure production time, and can be subdivided further into three main modes: the laser exposure production time, and can be subdivided further into three main modes: the laser exposure the feed containers by mode, the recoating mode and some other activities like the filling of the feed containers mode, the recoating mode and some other activities like an automatic powder supply system.

Based on 63 batches (5801 products), Figure 3 shows the cts), Figure 3 shows the an automatic powder supply system time distribution for products made of PA2200 with a layer thickness of 0.12 time distribution for products made of PA2200 with a layer thickness of12mm.PreheatingExposure59.40%27.48% 27.48%Recoating0.02%Other5.93% 5.93%1.24%5.93%Cooling Down Cooling DownCleaningFigure 3: Production modes (PA2200, 0.12mm) [11] Figure 3:Of course, the contribution of the productive modes depends on the product and batch Of course, the contribution of the productive modes depends on the product and batch Of course, the contribution of the productive modes depends on the product and batch design (e.g. total product volume and batch height). Based on multiple batches, an average total product volume and batch height). Based on multiple batches, an average total product volume and batch height).

Based on multiple batches, an average fferent layer thicknesses distribution of the productive modes is listed in Table 1 for two different layer thicknesses distribution of the productive modes is listed in Table 1 for two di m) and two materials (Fine Polyamide PA2200 and Glass-filled Fine (0.12mm and 0.15mm) and two materials (Fine Polyamide PA2200 and Glass m) and two materials (Fine Polyamide PA2200 and Glass Polyamide PA3200GF [9] ).EOSINT P760PA2200PA2200PA3200GFLayer thickness0.12mm0.15mm0.15mm#batches631211Laser Exposure31.45%48.71%21.95%Recoating68.53%51.25%78.02%Other Activities0.02%0.04%0.03%Standard Error1.13%5.97%2.52%Standard deviation (σ)8.97%20.68%8.35%Table 1: Distribution of the productive modes [11]Table 1 and Figure 3 show the significant contribution of the recoating mode to the total productive mode ranging from 50% up to 80% depending on the layer thickness and the material used.

Consequently, the part (batch) height has a major influence on the total production time.Multiple linear regression analyses indicate that the total production time (tprod) is a function of the total batch volume (V) and batch height (h). Equation 1 presents the formula for the first combination of material and layer thickness (PA2200 / 0.12mm).tprod [s] = 8250 + 2135 * h [cm] + 1.39 * V [cm³]with a R² = 0.96Since the product volume is determined by the product-design, the total production time can be minimized by optimizing the product orientation or nesting efficiency for batches with multiple products in order to reduce the batch height. These conclusions are consistent with previous work by Mognol et al. [12].Power StudyThe power levels of all different subunits during the productive modes of the EOSINT P760 are presented in Figure 4.

[12].Power StudyThe power levels of all different subunits during the productive modes of the EOSINT P760 are presented in Figure 4. The two most important energy consumers are the Laser Cooling Unit (3kW) and the Process Chamber Heaters (ranging from 0.8kW to 3.6kW in order to preheat the new powder material when a new layer is coated.). The peak consumption (up to 8.5kW) during the recoating process is mainly caused by the process chamber heaters warming up the new added powder layer.Similar power measurements were performed for EOSINST P360 (including an external Laser Cooler Unit of 2kW) and EOSINT FORMIGA P100 machine tools.

In comparison with the data available in literature (Table 3), the EOSINT P760 has a significant lower productivity. Nevertheless, the energy consumption per kilogram is comparable.EOSINT P760PA2200PA2200PA3200GFLayer thickness0.12mm0.15mm0.15mmProductivity (kg/h)0.170.150.24Energy / hour (kWh/h)6.205.976.32Energy / kg (kWh/kg)36.539.826.3Table 4: Process productivityConsumables StudyEOSINT machine tools, considered here, are equipped with compressed air driven nitrogen generators to create the inert atmosphere. The compressed air flow rates for the different machine tool types are listed in Table 5.Flow RateEOSINT P76020 m³/hEOSINT P3606 m³/hEOSINT FORMIGA P10010 m³/hTable 5: Compressed Air Flow Rates [9]Other SLS machine tools often use externally produced gasses (e.g.

nitrogen or argon)to create the required protective atmosphere.As indicated in the CO2PE!-methodology [7], the total material flow (determined by the product design and equal for all processes in a comparative study) is not relevant for a Unit Process Life Cycle Inventory (UPLCI), but the generated amount of waste is process dependent and included in our study.Despite the wide coverage of material production processes in LCI-databases, the production processes (e.g. atomisation processes) of powder materials used in SLS is still poorly documented in terms of their environmental impact.For SLS processes, Telenko et al. [16] approximated the specific energy consumption for 1 kilogram of nylon 12 powder. The final estimate for energy use during nylon manufacture is 85 MJ/kg using a mix of electricity and energy stored in fossil fuels, however neglecting the precipitation process [16,17,18].Based on industrial observations, half of the remainder powder of SLS is recycled.

Table 6 shows the powder utilization (PU) rates based on weight and volume as well as the waste flow rates for the different combinations of material and layer thicknesses.

The waste material rates around 45% fit well within the manufacturer’s recommendations about the refresh rates, which is the ratio between virgin and used powder (Table 7), using a very high portion of new powder material [9,15,18].EOSINT P760PA2200PA2200PA3200GFLayer thickness0.12mm0.15mm0.15mm#batches631211PUWeightVolume5.2%10.2%5.9%12.2%3.7%7.2%Waste material (Weight Based)44.9%43.9%46.4%Table 6: Powder utilization rates [11]MaterialRefresh rate (%) 30 to 50 50 to 70 30 + (30 overflow) 50 + (overflow)PA2200 PA3200GF DuraForm GF DuraForm Table 7: Recommended refresh rates [9,15]EOS GmbH [9]3D-Systems [15]Selective Laser Melting (SLM)Time studyFor SLM processes a time study is performed on a Concept Laser M3 Linear machine [10], using AISI 316L stainless steel (X5CrNi18-10) and a layer thickness of 30 µm.

Four use modes have been identified in accordance with the process principle: preheating the process chamber and creating the inert atmosphere (always 30 minutes), exposure (i.e. selective scanning by the laser), recoating a new powder layer, and final cooling down and part/powder extraction (cleaning).

Figure 5 shows the time share of the identified modes, based on a sample batch with a total production time of four hours.12%68%Preheating (30min) Preheating (30min)Exposure (Melting) Exposure (Melting)15%Recoating5%Cooling Down + Cleaning Cooling Down + CleaningFigure 5: Share of the different modes FigurePower studyDuring the machine tool analysis of the Concept Laser M3 Linear machine [10], 8 During the machine tool analysis of the Concept Laser M3 Linear machine [10], 8 During the machine tool analysis of the Concept Laser M3 Linear machine [10], 8 different subsystems were identified: laser unit (including the laser cooling unit), powder different subsystems were identified: laser unit (including the laser cooling unit), powder different subsystems were identified: laser unit (including the laser cooling unit), powder ng platform, coater, XY-Positioning of the scanner, nitrogen dosage chamber, building platform, coater, XY Positioning of the scanner, nitrogen During the productive modes circulation unit, cabinet cooling and the computer unit.

During the pro circulation unit, cabinet cooling and the computer unit. (exposure and recoating modes) Figure 6 shows the power levels for all subsystems during the productive modes. The peaks (around 3.5 kW) power levels for all subsystems during the productive modes. The peaks (around 3.5 kW) power levels for all subsystems during the productive modes. The peaks (around 3.5 kW) ) is moving. The laser unit coincide with periods during which the recoating system (sweeper) is moving coincide with periods during whi (2.24 kW for an output power of 100 W), is with 68% of the total machine tool power the (2.24 kW for an output power of 100 W), is with 68% of the total machine tool power the (2.24 kW for an output power of 100 W), is with 68% of the total machine tool power the most power consuming subunit. Table 8 lists the average machine tool power levels during all lists the average machine tool power levels during all most power consuming subunit.

Table 8 lists the average machine tool power levels during all lists the average machine tool power levels during all most power consuming subunit. Table modes.ting modes), the laser source is firing continuously. Figure) t t a W(r e w o P400035003000250020001500Ex Re Ex Re Ex Ex Re Ex Re ExMachine (Total) Machine (Total) Laser Unit (incl. Cooling) Laser Unit (incl.

Figure) t t a W(r e w o P400035003000250020001500Ex Re Ex Re Ex Ex Re Ex Re ExMachine (Total) Machine (Total) Laser Unit (incl. Cooling) Laser Unit (incl. Cooling) Dosage Chamber Dosage Chamber Building Platform Building Platform Sweeper Sweeper Compound table system Compound table system Computer Computer Cooling Cooling Nitrogen Circulation Nitrogen Circulation Other Units (Not plotted) Other Units (Not plotted)10005000: Power levels during productive modes of Concept Laser M3 Linear Figure 6: Power levels during productive modes of Concept Laser M3 Linear : Power levels during productive modes of Concept Laser M3 Linear(Ex = Exposure, Re = Recoating) [11] (Ex = Exposure, Re = Recoating)The energy consumption of another SLM machine tool: the EOSINT M250 Xtended The energy consumption of another SLM machine tool: the EOSINT M250 Xtended The energy consumption of another SLM machine tool: the EOSINT M250 Xtended ] was analysed by Mognol et al. [12].

[12]. The average electrical powers during standby and [9] was analysed by Mognol et al. [12]. The average electrical powers du ] was analysed by Mognol et al. [12]. The average electrical powers du operation mode were 2 kW and 4 kW respectively. operation mode were 2 kW and 4 kW respectively.Concept Laser M3 LinearPowerPreheating + Creating Inert Atmosphere2.25 kWExposure (Melting)3.25 kWRecoating3.45 kWCleaning + Product Removal0.7 kWTable 8: Average Power LevelsConsumables studyTo avoid oxidation problems, SLM processes operate in a protective (inert) atmosphere. Since nitrogen is used to create this protective atmosphere in the process chamber of the Concept Laser M3 Linear machine for stainless steel, this should be taken into account as consumable. A pre-flushing rate of 6.5 m³/h during the start-up mode (30 min) is followed by a continuous flow rate of 3.5 m³/h of nitrogen during the actual production phase.Morrow et al.

A pre-flushing rate of 6.5 m³/h during the start-up mode (30 min) is followed by a continuous flow rate of 3.5 m³/h of nitrogen during the actual production phase.Morrow et al. [19] indicate the required specific energy consumption (SEC) and generated airborne emissions for direct and indirect (where powder is produced from finished metal slabs or plates) H13 steel powder production processes, as listed in Table 9.(MJ/kg)Electricity Consumption Other Energy Consumption (e.g.

fuel oil, natural gas)Atomization Process Direct 15.90Indirect 26.861.725.95Table 9: Specific Energy Consumption (SEC) [19]It is observed that, relative to the production of steel plate (20.4 MJ/kg), the direct atomization route to producing steel powder requires approximately 20% less energy per kilogram of produced powder while the indirect atomization route to produce steel powder requires approximately 25% more energy [19].Based on five consecutive batches (with rather small parts) of the same material (X5CrNi18-10), including one complete emptying and cleaning operation of the building platform as well as the feed container, a ratio of 20.4% was found between the weight of the waste material and the weight of the product. This means that for a product of 1 kg, 204 g waste material is created (e.g.

This means that for a product of 1 kg, 204 g waste material is created (e.g. filter residues, aerosol emissions, post processing including the EDM cut to remove the parts from the base plate).Life Cycle Analysis (LCA)In this section, for both SLS and SLM, a Life Cycle Analysis (LCA) of a sample batch is described in order to identify the most important contributors to the environmental impact of the processes. The impacts are calculated based on the Europe ReCiPe H/A-method [20] using the ecoinvent database [21] and expressed in millipoints (Pts). Incineration is selected as end of life (EOL) treatment for the waste powder materials.

Incineration is selected as end of life (EOL) treatment for the waste powder materials. Process emissions (gaseous and aerosols) are not yet included.Selective Laser Sintering (SLS) Selective Laser Sintering (SLS)The SLS case study took place on an EOSINST P760 machine tool for a sample batch ook place on an EOSINST P760 machine tool for a sample batch ook place on an EOSINST P760 machine tool for a sample batch 32 kg and total production time of ith a total product volume of 3.32 kg and total production time of and Figure 7 show theof PA2200 material with a total product volume of 3. around 15 hours. The used layer thickness is 12 environmental impact and share of the different contributors respectively15 hours. The used layer thickness is 120µm.

Table 10 and Figureshare of the different contributors respectively.Impact (Pts)Energy Consumption Energy Consumption120 kWh6.34Compressed Air Consumption Compressed Air Consumption340 m³3.53Waste material(PA2200)Powder ProductionEnd of Life Treatment10.27 kg7.801.16Table 10: LCA of SLS sample batchEnergy Energy 33.7% 33.7%Compressed Air 18.7%Waste Material Production 41.4%Waste Material EOL 6.2%Figure 7: LCA of SLS sample batchSelective Laser Melting (SLM) Selective Laser Melting (SLM)iew of the environmental impact created during and Figure 8 show an overview of the environmental impact the production of a sample batch with a total product weight of 409 gram and a total the production of a sample batch with a total product weight of 409 gram and a total the production of a sample batch with a total product weight of 409 gram and a total production time of 4 hours on the Concept Laser M3 Linear machine tool.

of 4 hours on the Concept Laser M3 Linear machine tool.Table 11 and Figure 8Impact (Pts)Energy Consumption Energy Consumption11 kWh0.581Nitrogen Consumption Nitrogen Consumption15.5 m³0.743Waste material(X5CrNi18-10)Powder ProductionEnd of Life Treatment0.084 kg0.0420.002Table 11: LCA of SLM sample batch Table 11Energy Energy 51.9% 51.9%Nitrogen 45.3%Waste Material EOL 0.1%Waste Waste Material Material Production Production 2.6%Figure 8: LCA of SLM sample batchOf course, the relative proportions will change depending on the product (batch) size Of course, the relative proportions will change depending on the product (batch) size Of course, the relative proportions will change depending on the product (batch) sizeand design.Improvement PotentialTables 10 and 11 indicate that the environmental impacts are caused by several indicate that the environmental impacts are caused by several indicate that the environmental impacts are caused by several factors.

Reduction measures for all these impact creating factors are investigated and the tors. Reduction measures for all these impact creating factors are investigated and the tors. Reduction measures for all these impact creating factors are investigated and the provement potential for both SLS and SLM is described below. improvement potential for both SLS and SLMFirst, the total energy consumption can be minimized by selectively switching on and First, the total energy consumption can be minimized by selectively switching on and First, the total energy consumption can be minimized by selectively switching on and over the different modes. For example, the EOSINT P360 and P760 off the subsystems over the different modes.

For example, the EOSINT P360 and P760 off the subsystems over the different modes. For the EOSINT P360 and P760 laser cooling unit, which consume 2kW and 3kW respectively, which is machine tools have a laser cooling unit, which consume 2kW and 3kW respectively, laser cooling unit, which consume 2kW and 3kW respectively, , including the preheating and off-modes. Since identified production modes, including the preheating and off active during all identified production platform heaters...) are (e.g. process chamber heaters, frame heaters, platform the process heaters (e.g. process significant energy savings responsible for an important part of the total energy consumption, significant responsible for an important part of the total energy consumption, can be obtained by better thermal insulation.

can be obtained by better thermal insulation.Another reduction measure can be found in the introduction of external cooling down tion measure can be found in the introduction of external cooling down tion measure can be found in the introduction of external cooling down cycles. By use of a second process container, the total machine time can be reduced by 2 cycles. By use of a second process container, the total machine time can be reduced by 2 cycles.

By use of a second process container, the total machine time can be reduced by 2 cycles. By use of a second process container, the total machine time can be reduced by 2 cycles. By use of a second process container, the total machine time can be reduced by 2 hours (at standby power level consumption) and the residual heat of the previous build can be hours (at standby power level consumption) and the residual heat of the previous build can be hours (at standby power level consumption) and the residual heat of the previous build can be recovered partially.Furthermore, the total energy consumption depends mainly on the total production Furthermore, the total energy consumption depends mainly on the total production Furthermore, the total energy consumption depends mainly on the total production ated by Equation 1.

time, which is a function of the build volume and height as indicated by Equation 1 time, which is a function of the Consequently, the energy consumption can be reduced by minimizing the building height of Consequently, the energy consumption can be reduced by minimizing the Consequently, the energy consumption can be reduced by minimizing the the batch, for example by optimizing the part orientation or nesting efficiency of the batch. the batch, for example by optimizing the part orientation or nesting efficiency of the batch. the batch, for example by optimizing the part orientation or nesting efficiency of the batch.To create the required protective atmosphere in the process chamber, internally or To create the required protective atmosphere in the process chamber, internally or To create the required protective atmosphere in the process chamber, internally or externally generated nitrogen (or another inert gas) is often used.

In both cases a significant externally generated nitrogen (or another en used. In both cases a significant environmental impact is created by the consumption of compressed air or nitrogen from the environmental impact is created by the consumption of compressed air or nitrogen from the environmental impact is created by the consumption of compressed air or nitrogen from the external reservoir. Besides a reduction of the total production time (e.g. nesting efficiency, Besides a reduction of the total production time (e.g. nesting efficiency, Besides a reduction of the total production time (e.g. nesting efficiency, build height), another reduction m easure for this consumption can be found in a better sealed another reduction measure for this consumption can be found in a better sealed process chamber, so there is no longer need for continuous refilling of the process chamber. process chamber, so there is no longer need for continuous refilling of the process chamber.

process chamber, so there is no longer need for continuous refilling of the process chamber. process chamber, so there is no longer need for continuous refilling of the process chamber.Furthermore, and in particular for SLS, a large environmental impact is generated by Furthermore, and in particular for SLS, a large environmental impact is generated by Furthermore, and in particular for SLS, a large environmental impact is generated by the high amount of waste powder. Besides measures to increase the nesting efficiency (e.g. f waste powder. Besides measures to increase the nesting efficiency (e.g. f waste powder. Besides measures to increase the nesting efficiency (e.g.

f waste powder. Besides measures to increase the nesting efficiency (e.g. f waste powder. Besides measures to increase the nesting efficiency (e.g. reduction of the building height), more optimal recycling procedures can reduce the amount reduction of the building height), more optimal recycling procedures can reduce the amount reduction of the building height), more optimal recycling procedures can reduce the amount of waste and related environmental impact significantly. Dotchev et al. [19] provide a of waste and related environmental impact significantly. Dotchev et al. [19] provide a of waste and related environmental impact significantly. Dotchev et al. [19] provide a tical method to collect and separate un-sintered powder. In order to control the input practical method to collect and sintered powder.

Dotchev et al. [19] provide a tical method to collect and separate un-sintered powder. In order to control the input practical method to collect and sintered powder. In order to control the inputmaterial quality and use the fresh powder more efficiently, they suggest using different grades according to the melt flow rate of the recycled powder.Finally, the impact created by almost all factors mentioned above can be reduced by a well-considered choice of the SLM/SLS equipment (e.g. optimal size of the process chamber and build container). Flexible adaptable process chambers (height, width as well as depth) and building containers can allow producing a wide variety of products (batches) on the same machine tool, limiting the created environmental impact.Resource Efficiency of SLS on System LevelAmong different available approaches, the exergy metric for discrete manufacturing processes applications has been under development throughout the last decade [22,23].

Exergy is defined by Szargut et al. [24] as “the amount of work obtainable when some matter is brought to a state of thermodynamic equilibrium with the common components of the natural surroundings by means of reversible processes, involving interaction only with the abovementioned components of nature”. Unlike energy, exergy is not conserved; thus, exergy can be destructed in a system.As the exergy analysis in power generation and process industry systems, the exergy analysis in a discrete manufacturing system aims to detect and evaluate thermodynamic imperfection and indicate possibilities for improvement. The exergy efficiency is an important parameter in achieving these objectives. In the present study, the exergy analysis is performed for the SLS process.First the input-output streams data need to be gathered. The exergy value for different streams is taken from Szargut et al. [24], unless stated otherwise.

The exergy value for different streams is taken from Szargut et al. [24], unless stated otherwise. The specific chemical exergy value of PA2200 is based on Nylon 6 calculated by Dewulf et al. [25]. The compressed air input is fed into nitrogen generator inside the machine to produce nitrogen gas. The nitrogen is utilized to create an inert atmosphere inside the process chamber. The products, unused material and purged gases are the output of the SLS process.

The nitrogen is utilized to create an inert atmosphere inside the process chamber. The products, unused material and purged gases are the output of the SLS process. The summary of exergy analysis of the same build as in Table 10 is given in Table 12.InputOutputMaterialPowder (PA 2200) Compressed air Electricity Product Unused material Purged gas (N2) Purged gas (O2)Mass (kg)23.87 4083.33 20.54 90 318Specific Chemical Exergy (kJ/kg) 33000 15033000 33000 25.7 124Exergy (kJ)787710 61200 432000 109890 677820 2313 39432Total Exergy (kJ) 1280910829455Table 12: Exergy analysis of the SLS processFrom the data in Table 12, the exergy efficiency of the SLS process can be calculated accordingly. Here, the degree of perfection ηp will be used as the efficiency metric [23,24,26]. It is the ratio of the useful output exergy, in this case the product exergy, to the total input exergy. Based on this definition, the calculated value of ηp is approximately 8.6%.

It is the ratio of the useful output exergy, in this case the product exergy, to the total input exergy. Based on this definition, the calculated value of ηp is approximately 8.6%. It should be mentioned that this value disregard the possibility of material recycling, which will be treated below.clearly shows one of the advantages of exergy analysis compared to the more Table 12 clearly shows one of the advantages of exergy analysis compared to clearly shows one of the advantages of exergy analysis compared to conventional energy analysis, namely the inclusion of workpiece material and consumables in energy analysis, namely the inclusion of workpiece material and consumables in energy analysis, namely the inclusion of workpiece material and consumables in the analysis. This way, the improv ement potentials, which will be concluded from the the analysis.

This way, the improv ement potentials, which will be concluded from the the analysis. This way, the improvement potentials, which will be concluded from the analysis, will cover more domains than just energy efficiency. These other domains include analysis, will cover more domains than just energy efficiency. These other domains include analysis, will cover more domains than just energy efficiency. These other domains include process gasses and waste materials. An example of the improvement potential is the effects of process gasses and waste materials. An example of the improvement potential is the effects of process gasses and waste materials. An example of the improvement potential is the effects of material recycle in SLS process. SLS process. Recycling of powder material is a standard practice for SLS Recycling of powder material is a standard practice for SLS processes [18]. For the specific case in this study, it is reported that approximately 50% of the ].

For the specific case in this study, it is reported that approximately 50% of the ]. For the specific case in this study, it is reported that approximately 50% of the ]. For the specific case in this study, it is reported that approximately 50% of the remaining powder material can be recycled. An exergy diagram (Grassmann diagram) remaining powder material can be recycled. An exergy diagram (Grassmann diagram) remaining powder material can be recycled. An exergy diagram (Grassmann diagram) recycling is shown in Figure 9. By including illustrating the SLS process along with materia the recycled material, the degree of perfection ηp increases up to 12%. Care should be taken in the recycled material, the degree of perfection increases up to 12%. Care should be taken in determining the system boundary while calculating the system efficiency. In this case, in determining the system boundary while calculating the system efficiency.

Care should be taken in determining the system boundary while calculating the system efficiency. In this case, in determining the system boundary while calculating the system efficiency. In this case, in determining the system boundary while calculating the system efficiency. In this case, in order to show the positive effect of material recycling, only the fresh powder material is order to show the positive effect of material recycling, only the fresh powder material is order to show the positive effect of material recycling, only the fresh powder material is included as material input.strating the SLS process along with material recycling is shown in Figure 9Figure 9: Exergy diagram of the SLS process FigureIn Figure 9, the external exergy losses include the purged process gas and stray heat , the external exergy losses include the purged process gas and stray heat , the external exergy losses include the purged process gas and stray heat transfer to the surrounding, e.g.

during cooling phase. The waste material is also considered as transfer to the surrounding, e.g. during cooling phase. The waste material is also considered as transfer to the surrounding, e.g. during cooling phase. The waste material is also considered as external exergy loss. Another type of exergy losses is the internal exergy losses, which are nother type of exergy losses is the internal exergy losses, which are nother type of exergy losses is the internal exergy losses, which are process inherent losses and can be attributed to heat and friction losses. These internal and process inherent losses and can be attributed to heat and friction losses. These internal and process inherent losses and can be attributed to heat and friction losses. These internal and external losses provide two different kinds of improvement potentials.

These internal and process inherent losses and can be attributed to heat and friction losses. These internal and external losses provide two different kinds of improvement potentials. external losses provide two different kinds of improvement potentials.internal losses can be labeled process-improvement since these improvement since these measures will affect the internal process variables. For example, the application of more measures will affect the internal process variables. For example, the application of measures will affect the internal process variables. For example, the application of efficient laser sources will reduce the power consumption of the laser unit. This specific will reduce the power consumption of the laser unit. This specific will reduce the power consumption of the laser unit. This specific measure does not affect other variables outside the process itself. On the other hand, not affect other variables outside the process itself.

This specific measure does not affect other variables outside the process itself. On the other hand, not affect other variables outside the process itself. On the other hand, not affect other variables outside the process itself. On the other hand, improvement measures over external exergy losses can be originated from the modification of improvement measures over external exergy losses can be originated from the modification of improvement measures over external exergy losses can be originated from the modification of variables outside the process; hence the name systems-improvement. A better nesting improvement. A better nesting variables outside the process; hence the name systems efficiency is one example of this type of improvement. The nesting efficiency itself is a iency is one example of this type of improvement. The nesting efficiency itself is a iency is one example of this type of improvement.

The nesting efficiency itself is a iency is one example of this type of improvement. The nesting efficiency itself is a iency is one example of this type of improvement. The nesting efficiency itself is a process planning variable, not a process variable. Another example is the re-utilization of the process planning variable, not a process variable. Another example is the re process planning variable, not a process variable. Another example is the reThe effort to reduce internal losses can be labeled processwaste material. As proposed by Dotchev et al. [19], the degraded powder material can be reused for a trial work, for example. This measure might not increase the resource efficiency of the process but it improves the overall efficiency in the broader spatial and longer temporal scale.Therefore, the exergy analysis provides a useful view on the resource efficiency of the SLS and SLM processes at system level.

While typical energy analysis manages to illustrate different power flows inside the system, leading to a better focus on proposing energy efficiency improvement measures, the exergy analysis delineates prospective improvement potential not only from energy point of view, but also from material and consumables utilization. It also pinpoints improvement potentials which are related to systems variables, such as nesting efficiency and material recycling. Furthermore, the exergy efficiency serves as a convenient resource efficiency metric which combines the energy and material utilization.SummaryUsing the CO2PE!-Methodology [4-5], Life Cycle Inventory (LCI) data is collected and the environmental performance regarding energy and process gas consumption as well as generated waste material of both SLM and SLS processes is investigated in this paper. Impact reducing measures are presented for the most important contributors to the environmental impact of both processes.

Impact reducing measures are presented for the most important contributors to the environmental impact of both processes. Finally the resource efficiency of SLS processes is presented using the exergy metric. Future work will include the completion of the LCI data regarding process emissions and powder production processes.AcknowledgementsWe acknowledge the support of the Institute for the Promotion of Innovation through Science and Technology in Flanders (IWT-Vlaanderen) through its PhD grant N°091232, and of the European Fund for Regional Development (EFRO - Europees Fonds voor Regionale Ontwikkeling) and the Agentschap Ondernemen (Flemish government) through the D2 project 476.One of the co-authors, Evren Yasa, thanks TUBITAK (The Scientific and Technological Research Council of Turkey) for its financial support given under “Ph.D.

support program for students in foreign countries” and acknowledges K.U.Leuven for the support through the project GOA/2002-06 and IWT for the support through the SBO-project DiRaMaP.References[1] Over, C., Meiners, W., Wissenbach, K., Lindemann, M., Hutfless, J., 2002, Laser Melting: A New Approach for the Direct Manufacturing of Metal Parts and Tools, Proc. Euro-uRapid, Int. Users Conf., A-5.[2] Kruth, J-P., Vandenbroucke, B., Van Vaerenbergh, J., Mercelis, P., 2005, Benchmarking of different SLS/SLM processes as rapid manufacturing techniques, Proceedings of 1st International Conference of Polymers and Moulds Innovations, Gent, Belgium.[3] Kruth, J.-P., Levy, G., Klocke, F., Childs, T.H.C., 2007, Consolidation phenomena in laser and powder-bed based layered manufacturing, CIRP Annals – Manufacturing Technology 56(2), pp.

 Energy Dissipative Constitutive Model for Multi-Surface Interfaces at Weld DefectSites in Ultrasonic ConsolidationNachiket Patil, Deepankar Pal and Brent E. Stucker Industrial Engineering, University of Louisville, KY 40292A new finite element based constitutive model has been developed for quantification of energy dissipation due to friction and plastic deformation at the mating interface of two surfaces during the Ultrasonic Consolidation process. This work will include bridging the mesoscopic response of a dislocation density based crystal plasticity finite element framework at inter and intra-granular scales and a point at the macroscopic scale. This response will be used to develop an energy dissipative constitutive model for multi-surface interfaces at the macroscopic scale. The constitutive model will be used for quantification of energy consumed at lack of fusion and trapped oxide defects present in the build and the amount of energy input required to compensate for it.

This numerical procedure will help in real time optimization of process parameters and closed loop control.IntroductionMetal based additive manufacturing processes include both thermally-induced fusion processes and solid state processes. Some of the benefits of a solid state process lie in unique microstructures made possible because of near room temperature processing and the avoidance of solidification cooling induced residual stresses in the built part. These benefits can be further exploited if closed loop control of these processes becomes possible. The major hurdle in this is the computational complexity of the contact simulations, especially at the defect sites and the layer interfaces, involved in solid state processes like Ultrasonic Consolidation (UC).

The present work is focused on development of a novel framework for interface simulations with the capability to include friction and various large deformation, nonlocal and geometrical nonlinearities involved in metal deformation.Ultrasonic ConsolidationUltrasonic consolidation (UC) is a solid-state additive manufacturing process which combines ultrasonic metal welding and milling to produce three dimensional objects. The process uses the high frequency ultrasonic shearing vibration at low amplitude and a normal force to break the oxide layers between two foils and then bond them by bringing together two nascent foil surfaces. Figure 1 shows a schematic of the UC process. The weld quality achieved in the process is sensitive to machine parameters. This sensitivity calls for better optimization of machine parameters through well informed computer simulations.The UC process involves nonlinear plasticity which is inherently multi scale in space as well as time.

The nonlinear and nonlocal plasticity in the process is dependent on the rate of deformation as well as size effects. The phenomenon of acoustic softening during UC is governed by these complicated material behaviors along with rate dependent dislocation dynamics. Continuum plasticity can be used to model the average phenomenon involved but it cannot account for the grain fragmentations or grain structure evolutions, which are at the core of the process if the quality of the weld is in question. A computational framework involving dislocation density based crystalplasticity [1,2] has been developed.

A computational framework involving dislocation density based crystalplasticity [1,2] has been developed. The computational cost involved in these simulations can be high and thus further work is required to reduce it by an order of magnitude to make it fast enough to work with machines for online process control.In the present work a novel framework for developing interfacial constitutive models is presented to evaluate the macroscopic response of interfaces for various surface roughness values and material properties along with homogenization of dislocation dynamics near interfacial regions.Figure 1: Schematic of UC process [1]Interfacial Constitutive ModelThe concept of an interfacial constitutive model has been researched for various applications like crack growth and propagation. The cohesive zone model [3,4] is considered as one of the pioneering interfacial constitutive models.

The cohesive zone model [3,4] is considered as one of the pioneering interfacial constitutive models. The concept of a cohesive zone model can be further improved to simulate continuum level constitutive modeling of interfacial friction and sliding behavior [5].Interfacial constitutive model literature is discussed in this subsection because it is closely related to the new formulation attempted in this work. The Extended Finite Element (XFEM) [6] methodology recently employed for multi-material and discontinuity simulations is another important development which has the capability to deal with various jump conditions involved at material interfaces. XFEM deals with discontinuities by decoupling a macroscopic average problem from the discontinuity problem at the FEM point integration level in order to calculate its effect on the stiffness matrices and macroscopic unbalanced forces.

The decoupling of two problems rely on the functional evaluation of the coupling between the average macroscopic problem and the discontinuous behavior which could be a multi-scale phenomenon.Elastic plastic interfacial constitutive modeling Literature [5]Figure 2: Interface between two bodies [5]A cohesive interface between two bodies as shown in figure 2 can be describes as a contact interface between two bodies with or without any bonding mechanisms present at the interface. A constitutive model [5] for such an interface was developed by Su et. al and is discussed here to consider its potential application to Ultrasonic Consolidation. Consider two bodies (cid:2010)(cid:2878) and (cid:2010)(cid:2879) separated by an interfaceℑ. Assume {(cid:2187)(cid:3548) (cid:2778), (cid:2187)(cid:3548) (cid:2779), (cid:2187)(cid:3548) (cid:2780)} is an orthonormal triad.

Finally the required consistency condition for plastic deformation is as follows.(cid:1874)(cid:3036)∅(cid:3036) = 0, when ∅(cid:3036) = 0(4)(5)(6)(7)(8)The consistency equation helps identify the possibility of plastic deformation. The magnitude of plastic deformation has to be further determined based on the evolution of (cid:1871)(cid:3036) as a function of time. The evolution mechanisms for (cid:1871)(cid:3036) are in general described as hardening and softening and their evolution is a function of dislocation dynamics; and a detailed crystal plasticity simulation is required in order to capture them for general loading and geometry scenarios with given macrostructure information on both sides of the interface.Applications of the interfacial constitutive modelsThe interfacial constitutive law will help in simulation of the process at the macroscopic scale, which will result in reduced computational complexity.

Other applications for such a model include engineering problems such as cohesive modeling of crack interfaces during their propagation. The contact problem is also significant from the perspective of bonded or unbonded joints in metals. These energy dissipation laws will be also helpful in developing constitutive models and new interfacial elements for macroscopic vibration simulations.The application of such an interfacial model in additive manufacturing other than UC process will include support snapping simulation for Selective Laser Melting (SLM) process where an initial crack in the lattice will grow and cause support structures failure. In case of SLM delamination possibility of two successive layers built with significant recoat time can be also studied in depth.

In case of SLM delamination possibility of two successive layers built with significant recoat time can be also studied in depth. The friction stir surfacing process can be also modelled efficiently with the proposed simulation framework.Mathematical FormulationFinite Element FormulationThe schematic problem description is given in figure 4. Simple 3D brick elements are used for FEM simulation.Figure 4: Schematic diagram presenting the concept of an interfacial constitutive model for a macroscopic interface contactThe virtual work equation is written in virtual work formulation as(cid:2012)(cid:1848) = (cid:3505)((cid:2012)(cid:2013))(cid:1829)	((cid:2013))	(cid:1856)(cid:1874) + (cid:2012)(cid:1849)((cid:2012)(cid:1873)) = 0(9)(cid:2013) = [(cid:1828)](cid:2199)(10)The (cid:2199) vector is the nodal displacement vector.

Each split row will be added to the master Degree of Freedom (DOF) row based on the (cid:1840)(cid:3036) in the constraint equation for the master DOF. This methodology is explained in figure 5.Figure 5: Schematic diagram representing a slave node constraint on a contact surface and its solution methodology by splitting slave node stiffness as per the constraint equation.2.0 Free NodesA DOF is set free if it is sliding or has lost contact with the bottom surface.3.0 ((cid:2189)((cid:1876), (cid:1877)) ≤ 0) conditionTo simplify the algorithm, top surface nodes that are impinging are first allowed to impinge and then are forced back to the bottom surface iteratively.Contact Simulation AlgorithmThe contact simulation algorithm is based on minimizing unbalanced force iteratively along with updated contact constraints. First contact constraints are updated which will induce some unbalance forces due to shear slip and change in stiffness.

First contact constraints are updated which will induce some unbalance forces due to shear slip and change in stiffness. These unbalanced forces are solved in the new configuration which can further change the gap function and can change the contact constraints. The algorithm is presented in a flow chart format in figure 6.Figure 6: Flow chart showing FEM algorithm used for contact simulationCase StudyA simple case of surface to surface contact is considered here. Two prismatic blocks of size 0.001 × 0.001 × 0.0005	(cid:1865)(cid:2871) are in contact with each other as shown in figure 7. The z-direction height of each block is 0.0005m. The boundary conditions include a bottom surface fixed in all degrees of freedom and a top surface with slip and normal compressive displacement boundary conditions as shown in figure 7.

The correct boundary conditions for extracting constitutive models are outside the focus of this work, as this work focused on formulating a FEM framework which can consider different boundary conditions and give high fidelity contact simulations. The parameters of interest in this case study are the macroscopic aggregate shear resistance and macroscopic strain.The shear resistance of the contact surface is defined below.

The reason behind the fall of the shear resistance with increase in slip is due to contact being lost at the extreme left or right points due to the overhanging nodes at the end. Figure 9 shows a plot of macroscopic friction coefficient calculated as a ratio of S and N at each time instant during slip loading in the simulation. The reason for a macroscopic friction coefficient less than 0.3 of that of the actual interface can be attributed to the fact that the shear resistance at each point in the cross-section is not utilizedcompletely in one particular configuration or point in the time.

Another way to explain it is that the 0.3 is the maximum upper bound that can be achieved.The shear resistance will be totally lost at the overhangs because there is no material available on either side for contact friction as seen in the deformed shape shown in figure 10.Figure 9: Plot of the macroscopic friction coefficientConclusions and future workA novel framework for an interfacial friction model has been developed in the present work. The developed contact simulation framework shows good behavior at the contact interface.

The developed contact simulation framework shows good behavior at the contact interface. The novel features of this framework include:Mathematical framework for efficient assembly of FEM matrices  Solution of constraint equations exactly without any additional computational cost  No addition of penalty springs which can introduce spurious stiffening behavior at interfaceThe future work in this problem will include addition of the following features to the present framework in order to develop interfacial constitutive models or laws:Nonlinear simulations with crystal plasticity based material constitutive models for metals.   Periodic boundary condition considering the case of deriving a constitutive model for a point on an interfacial surface. Inclusion of surface roughnessFigure 10: Deformed and un-deformed mesh at a macroscopic shear strain of 0.05AcknowledgementsThe authors would like to acknowledge the Ofﬁce of Naval Research (ONR) for support through Grant No.

ergy Efﬁcient Modeling and Optimization of Additive Manufacturing ProcessesAnoop Verma, and Rahul Rai∗DART LAB, Mechanical and Aerospace Engineering Department,University at Buffalo, Buffalo-14260email:anoopver@buffalo.edu, rahulrai@buffalo.eduAbstractAdditive manufacturing (AM) is a leading technology in various industries including medical and aerospace for prototype and functional part fabrication. Despite being environmentally conscious, avenues pertaining to further reducing the impact of AM on the environment exist. Material wastage and energy consumption are two major concerns of the process that requires immediate attention. In this research, a multi-step optimization enabling additive manufacturing process towards energy efﬁciency is developed. Process objectives such as material waste and energy consumption are minimized both in part and layer domain. Numerous examples are presented to demonstrate the applicability of the developed approach.

Numerous examples are presented to demonstrate the applicability of the developed approach. The models formulated here for selective laser sintering (SLS) process can be easily extended to other additive manufacturing technologies.1 IntroductionAdditive manufacturing is a process of fabricating three-dimensional solid objects from a digital model such as CAD model. It is achieved using additive processes, where laying down successive layers of material creates an object. Additive manufacturing is distinct from traditional techniques (i.e. subtractive processes), which mostly rely on the removal of material by means of drilling, cutting etc. Over the last two decades, there have been tremendous technical advances in the ﬁeld of additive manufacturing, including input ﬁle formats [1, 2, 3], materials selection [4], and process planning [5, 6].Process sustainability is an important aspect which has received little attention in additive manufacturing.

Despite the noticeable advances in the additive manufacturing technologies over traditional machining, relatively very few studies have focused on studying its potential effect on the environment and energy efﬁciency [7, 8, 9]. With the recent trend towards green economy, research in additive manufacturing should be complemented with the energy efﬁciency and sustainability. In additive manufacturing, process sustainability indicators such as material waste and energy consumption are always a concern.For example, FDM processes generate waste in the form of support structure. Laser based processes such as SLA and SLS selectively sinters materials from a large material bed, leaving majority of un-sintered (56%) material as a waste [7, 10].

Laser based processes such as SLA and SLS selectively sinters materials from a large material bed, leaving majority of un-sintered (56%) material as a waste [7, 10]. On the other hand, AM sub-processes such as laser scanning, raw material deposition, material extrusion, and bed heating and cooling processes are energy intensive [11, 12].The research presented in the paper addressed the current research need towards sustainable additive manu- facturing. More speciﬁcally, process energy and material waste in AM process are minimized using a multi-step optimization approach. The contribution of proposed paper is two-fold namely, (1) process sustainability model- ing and optimization, (2) development of a novel heuristic optimization based adaptive slicing approach.

The contribution of proposed paper is two-fold namely, (1) process sustainability model- ing and optimization, (2) development of a novel heuristic optimization based adaptive slicing approach. Unlike other adaptive slicing approaches in the literature, the developed approach is generic and does not require pre- calculation of part geometry information such as complexity, curvature, and feature identiﬁcation [13, 14, 15].∗Corresponding AuthorThe paper is organized as follows. In Section 2, numerous objectives and the concerned optimization models are formulated. The overall solution methodology is demonstrated in Section 3. Experimental results and ﬁndings from the present analysis are discussed in Section 4.

The overall solution methodology is demonstrated in Section 3. Experimental results and ﬁndings from the present analysis are discussed in Section 4. Section 5 concludes the present work with some notes on future research directions.2 Mathematical formulation of the processThe energy consumed by the selective laser sintering (SLS) machine during whole part build can be categorized into: (1) processing energy, and (2) non-processing energy (Figure 1). Non-processing energy consists of energy consumed by machine during recoater arm movement, build piston movement (upward/downward), and initial heating. On the other hand, processing energy consumption is largely determined by the amount of material which is required to be fused together to build the whole part. Experimental studies in the literature have found out the contribution of processing energy as much as 56% of the total energy consumption.

Experimental studies in the literature have found out the contribution of processing energy as much as 56% of the total energy consumption. In the present research, both processing and non-processing energy consumption is minimized by (1) reducing the number of recoater arm movement, build piston movement, and (3) optimizing the laser scan movement to sinter the part for each layer. A set of underlying objectives are formulated next.Figure 1: Schematic of selective laser sintering (SLS) with highlighted energy components2.1 Laser proecssing energy The material absorptivity (ω), laser average intensity (Iavg), laser scan speed (SS), and laser spot diameter (SD) are the factors contributing to the overall input energy requirement (e) for sintering the material powder [13].

Mathematically, it can be written as:e =SD × ω × Iavg SSLaser average intensity (Iavg) is the laser power per unit area, written as:Iavg = LP/ASDIn equation 2, LP is the laser power, and ASD (equation 3) is the area covered by laser beam spot.ASD =π × SD2 4The equation (1) can thus be written as:e =4 × ω × LP π × SS × SD(1)(2)(3)(4)The total energy required by the laser to sinter one slice of the part will be:Es = e × AsIn equation (5), As is the area of sth slice of the part. As can be broken down to the actual number of scan lines and slice boundaries to sinter the power material. Figure 2 explains the effective slice area of a sample 2D proﬁle. It consists of scanning the outer boundary and inside solid area through a series of parallel laser beams. The slice sintering area can be written as:As = SD ×Is(cid:88)SVi +Cs(cid:88)Pji=1j=1In equation (6), As is the sintering area of sth slice proﬁle, SD is the beam spot diameter.

The slice sintering area can be written as:As = SD ×Is(cid:88)SVi +Cs(cid:88)Pji=1j=1In equation (6), As is the sintering area of sth slice proﬁle, SD is the beam spot diameter. SV is the length of scan vector, Is being the total number of scan vectors in a slice s. P is the outside boundary of slice, and Cs is the number of contours in slice s. Hence, the total energy required to sinter the whole part will be:Ξ =S (cid:88)s=1Es (cid:44) Ξ = e ×S (cid:88)s=1As (cid:44) Ξ =4 × ω × LP π × SS × SD×S (cid:88)s=1Is(cid:88)i=1SVsi +Cs(cid:88)j=1PsjIn equation (7), S is the total number of slices in part (i.e. equivalent to L-1, where L is the number of layers).Figure 2: Part slice component determining effective sintering area2.2 Material waste In SLS, the un-sintered power material is typically reused, however, at the expense of powder material strength.

The powder binding strength deteriorates at each reuse run, making around 56% of un-sintered material as waste. In the present research, the material wastage (ξ) is calculated at part level, mathematically can be written as:ξ = (SVbed − Vpart) × ψIn equation (8), Vpart is the volume of the 3D part, ψ is the material waste fraction, and SVbed is effective sintering volume of the bed. The powder material to be laid on sintering zone depends on the part height, thus the effective volume of the bed becomes ARbed × Hpart. The above equation is reduced to:ξ = ((ARbed × Hpart) − Vpart) × ψ(5)(6)(7)(8)(9)2.3 Surface quality The part surface ﬁnish is another important factor that needs to be considered while minimizing material waste. As mentioned in earlier sections, part height plays an important role in determining the energy consumption of the fabrication process.

As mentioned in earlier sections, part height plays an important role in determining the energy consumption of the fabrication process. Maintaining an appropriate surface ﬁnish may increase the process energy consumption, but will minimize part post-processing. Here, the part surface roughness is used as an estimate to calculate surface ﬁnish of the part based on the orientation of triangular facets. Mathematically, the surface roughness of a facet can be calculated as [16]: (cid:40)Rai =(2cosφsinφ × 937 + 3.5φ + 48) × 0.0254, if 0 <= φ <= π/2 (2cos(π − φ)sin(π − φ) × 937 + 3.5φ + 48) × 0.0254, otherwiseIn equation (10), Rai is the surface roughness of ith facet, φ is the angle between facets surface normal and build direction. Comparison of surface roughness of various facets is shown in Figure 3. The inclined facets will contribute to the surface roughness of the parts.

Comparison of surface roughness of various facets is shown in Figure 3. The inclined facets will contribute to the surface roughness of the parts. The average surface roughness is calculated as:Raavg =(cid:80)Fi=1 Rai × Ai (cid:80)F i=1 AiIn equation (11), Ai is the Area of ith facet and F is the total number of facets. In next section, the developed optimization models are discussed.Figure 3: Comparison of surface roughness of facets. (a) Inclined facets will add to the surface roughness of the part. (b)Parallel and perpendicular facets will have 0 surface roughness.2.4 Optimization models In this sub-section, description of the optimization models is provided. Objectives such as part surface roughness, volumetric error are combined with material waste and energy related objectives respectively to ensure the surface accuracy of the part is not sacriﬁced while part processing energy and material waste are minimized.

Considering the nature of addive manufacturing processes, the 3D models to be fabricated are optimized at two levels. In level 1, the 3D models as a whole is optimized by minimizing the material waste and surface roughness, whereas, in level 2, optimization is done for each slice of the part (to ensure minimum energy consumption) which is combined with the developed adaptive slicing approch.Overall framework of the two-level optimization approach is provided in Figure 4. The developed appriach starts with part preprocessing, where, 3D models are ﬁrst translated into positive x, y, and z axis for ease of calculation (step 1). Next, the part surface roughness and overall material waste is minimized (step2). It is achieved by optimallly orienting the part along build direction (i.e. z axis). The output of the level 1 optimization provides optimal orientation of the part which is kept ﬁxed.

It is achieved by optimallly orienting the part along build direction (i.e. z axis). The output of the level 1 optimization provides optimal orientation of the part which is kept ﬁxed. For level 2 optimization, both part slicing and the laser processing energy of individual slices is determined. Part sliicing is based on developed adaptive slicing heuristic which starts with random slicing solutions and iteratively improves the overall slicing while minimizing the process energy, and volumteric error of the part. Detailed description of the developed heurtistic algorithm is provided in Section 3.(10)(11)Figure 4: The developed optimization framework for adaptive slicing2.4.1 Minimization of material waste and part surface roughness ( level 1) In the part level, both material wastage and surface roughness are optimized.

The overall optimization model can be written as:Min.f (α, β) =w1 × ξ + w2 × Raavg w1 + w2, s.t., 0 <= α, β <= πIn equation (12), α, β, are the part rotation angles in x,y directions which are varied in the range (0 − π), w1, and w2 are the weights associated with the objective functions namely material waste and surface roughness. As the main objective of present study is to minimize material waste and energy consumption, higher weight is assigned to the material waste. In the research the weight associated with material waste crietia is set to 0.75 (i.e. w1=0.75), whereas, weight value of 0.25 (i.e. w2=0.25) is assigned to part surface roughness.2.4.2 Minimization of laser processing energy (level 2) The laser processing energy is optimized by varying the number of layers and scanning directions of the lasers for each layer. AM processes fabricate complex parts which often has varying cross sections.

AM processes fabricate complex parts which often has varying cross sections. The number of scan lines to sinter the 2D layer contour will vary from one direction to another. By optimizing the laser scanning directions, the laser processing energy can be minimized. Mathematically, it can be written as:Min.f (γ) =4 × LP π × SS × SD×I γ s(cid:88)i=1SVsi +Cs(cid:88)j=1Psj, s.t.0 <= γ <= π/2In equation (13), s is the indices used for slice, I γ s is the number of scan vectors required to be sintering for whole slice, γ being the laser scanning direction. The process energy consumption of each layer is minimized by optimizing the laser scanning movement.2.4.3 Minimization of overall energy and part volumetric error ( level 2) The overall energy consumption (Ξ) of the process is the laser processing energy (given in equation 13), summed up for all the slices of the part.

The adaptive slicing procedure selects the layer thickness so that the upper and lower contours of a slice have minimum difference. Volumetric error is used as an additional objective to ensure the parts required minimum post processing (another source of energy consumption). Volumetric error (VE) is deﬁned as the volume difference between the actual part and the fabricated part. Mathematically, it can be written(12)(13)as [8]:V E =L−1 (cid:88)ltl × |Al+1 − Al|l=1In equation (14), VE is the volumteric error, which is calculated using the area difference between consective layers, and slice thickness (lt). The overall objective function is written as:Min.f (lt, γ) =w1 × Ξ + w2 × V E w1 + w2, s.t., ltmin <= lt <= ltmaxIn equation (15), lt is the layer thickness which should be kept in the range ltmin − ltmax (i.e.

laser sintering capability), w1 and w2 are the weight associate with objective namely minimization of energy consumption and minimization of volumetric error. Higher weight is assigned to energy minimization objective (i.e. w1=0.75, w2=0.25). In next section, the methodology to solve the formulated objectives are discussed.3 Solution approachThe models developed in previous section level are optimized in two-level. In level 1, material waste and average surface roughness of part is minimized by optimally orienting the part towards build direction. Being highly non-linear, the objectives are optimized using a heurtic approach. The developed heuristic comprises of a local search technique namely simulated annealing (SA), and a genetic algorithm (GA) based meta-heuristic. Both the SA, and GA are well-known algorithms to solve highlly non-linear optimization problems as formuated in this paper.

Both the SA, and GA are well-known algorithms to solve highlly non-linear optimization problems as formuated in this paper. The developed heuristic unqiuely combines the traits of both SA, and GA to solve the underlying problem. SA is a point based evolutionary local search technique that explores the search space iteratively based on a probability distribution proportion to the temperature. In the developed research, SA explores the search space until no change in the objective function is found for consecutive 100 iterations [17]. Adaptive slicing is performed using genetic algorithm (GAs). GA’s belongs to a class of evolutionary search algorihms that works on population (also known as inital feasible solution) [18, 19]. GAs iteratively improves the search space by means of variation operators namely crossover, mutation, and selection mechanism.

GAs iteratively improves the search space by means of variation operators namely crossover, mutation, and selection mechanism. In the proposed research, individual refers to a set of layer thickness values constituting a part and a population consists of a set of distinct individuals. The cumulative layer thickness should be equal to the part height, which determines the number of layers. The process capability of the SLS machine is taken into consideration for population initialization and other processing steps. SA ﬁrst iteratively improves the objective function given in equation 14 and provides theFigure 5: Candidate solution generation and variation operators on sample 3D part(14)(15)Table 1: GA speciﬁc parameters used in the study Parameter No. Population size 1 Crossover type 2 Mutation type 3 Crossover probability 4 5 Mutation probability 6Value/range 10 single point swap 0.8 0.01Termination crietria Max.

Both SA, and GA are then applied in tandem to perform slice level energy minimization and adaptive slicing respectively. The quick convergence properties of SA provides faster solution, whereas, unique encoding schema of GA facilitates efﬁcient adaptive slicing. Figure 5 display the schamatic of GA encoding schema once the part is optimally oriented. The contribution of SA based search technique is also highlighted for random part slices. The developed developed candidate solution generation scheme ensures only feasible solutions are generated by: (1) generating part slice thickness within the range of machine capability, i.e. ltmin − ltmax, and (2) making the cumlative layer thickness of individual solution equal to the part height. Feasibility is also ensures after variation operators namely crossover and mutation is applied. The process shown in Figure 5 is repeated for each generation until the termination criteria is reached.

The process shown in Figure 5 is repeated for each generation until the termination criteria is reached. Here the termination criterion is maximum number of generations which is ﬁxed to 50. Genetic Algorithms (GA) is explore the search space to get the best solution in terms of slice thickness values and minimum energy consumption [20]. The GA speciﬁc parameters used in the study are shown in Table 1.In next section, details of experiment results is presented.4 Experimental resultsIn this section, numerous example parts are optimized using the solution methodology discussed in Section 3. The analysis shown here is based on SLS vanguard HiQ Sinterstation machine with polycarbonate (PC) as input material powder. Table 2 describes the machine speciﬁc parameters used in the present study1. In the present study, the laser processing energy is minimized, whereas, the non-processing energy used during piston, recoater arm movement, and bed heating is kept ﬁxed.

In the present study, the laser processing energy is minimized, whereas, the non-processing energy used during piston, recoater arm movement, and bed heating is kept ﬁxed. An interface is developed in MATLAB (2012a) for the analysis. Based on the formulated optimization models, the interface consists of two major components, i.e. part level optimizer, and slice level optimizer, discussed below. All example parts‘ used in the study are in mm.1http://www.3dsystems.com/products/dataﬁles/4.1 Part level optimizer Part level optimizer provides quick optimal solution to the material waste and surface roughness by appropriately orienting the part along build direction. An example 3D part of dimension (0.3mm × 0.3mm × 0.6mm) as shown in Figure 6 (a) is used for the analysis. The part level optimizer recommends the input part to be oriented 90◦, 0◦ along x and y axis respectively. The resultant surface roughness is 8.39 Ra, whereas, material wastage is 22.77 cm3 (Figure 6 (b)).

The resultant surface roughness is 8.39 Ra, whereas, material wastage is 22.77 cm3 (Figure 6 (b)). Material wastage reduced from 44.93 cm3 to 22.77cm3 (an improvement of 50.67%) (Figure 6 (c)).Figure 6: (a) an example 3D model, (b) the corresponding part level optimizer with inital and optimal part orien- tation using SA, and (c) iterative improvement in the material waste.4.2 Slice level optimizer The slice level optimizer performs optimization on part slices and provides optimal/near-optimal laser scan move- ment of each slice, slice thickness, the process build time, process energy consumption, and volumetric error. An input 3D part of dimension 2 mm × 2 mm × 4 mm (Figure 7(a)). The part level optimizer recommends the input part to be oriented as it is, i.e. 0◦, 0◦ along x and y axis respectively (Figure 7(b)-top right). The resultant surface roughness is 0.1438 Ra, whereas, material wastage is 299.87 cm3. A majority of material waste is attributed to the part proﬁle.

The resultant surface roughness is 0.1438 Ra, whereas, material wastage is 299.87 cm3. A majority of material waste is attributed to the part proﬁle. The process energy consumption is 0.79 W, whereas the process would take just 46 sec to build the part. Part volumetric error is 0.0635 mm3, suggesting high accuracy of the part. Figure 7 (b) also display the laser scanning movement of 6 random slices. Figure 7 (c) shows the energy consumption and build time of each slice, which follows similar trend, suggesting high correlation. .Figure 7: (a) an example 3D model, (b) the corresponding slice level optimizer with 6 random laser slices, and (c) the optimized energy consumption of each slice and corresponding build time.4.3 Results comparison In this section, the results obtained from the developed approach are compared against the state-of-the-art (i.e. uniform slicing).

uniform slicing). The models used for the comparison are actual models developed using additve manufacturingFigure 8: Comparison of proposed approach with state-of-the-arttechniques2. The parts includes here comprises of a variety of application areas with varying complexity, estimated by number of facets. Irrespective of the size, parts with large number of facets are considered to be more complex than parts with small number of facets. Depending on SLS process compatibility, the example parts are compared against uniform slicing with maximum layer thickness (0.1 mm), uniform slicing with minimum layer thickness (0.03 mm), and uniform slicing with mean layer thickness (0.065 mm).Figure 8 describes the list of 3D parts used in the analysis and the obtained results. Parts fabricated with maximum layer thickness are always quicker to build, and consumes least process energy. However, parts build using maximum layer thickness always yields poor surface quality (i.e.

However, parts build using maximum layer thickness always yields poor surface quality (i.e. higher volumteric error), resultantly, more energy consumption during post-processing. Compared with uniform slicing (with minimum, and mean layer thickness), proposed approach always provides the best results. In terms of material savings and surface roughness, proposed approach always provides best results. The savings in the material is larger for complex parts (i.e. parts with larger facets).2http://www.thingiverse.com/5 Conclusions and future researchThe experimental analyses demonstrate the applicability of optimization in additive manufacturing industry in achieving optimal process plan. The propsoed appoach outperforms state-of-the-art slicing approaches both in terms of material waste and energy consumption.

The propsoed appoach outperforms state-of-the-art slicing approaches both in terms of material waste and energy consumption. With an aim towards sustainability, the developed optimization models are most applicable as compared with state-of-the-art additive manufacturing.Considering the computational effort, only small-scale parts are analyzed for process optimization. Future research will aim towards minimizing the computational cost of optimization algorithms by developing process plans based on shape similairy. Modifying the developed approach to incorporate the process planning of other additive manufacturing technologies such as fused deposition modeling (FDM), selective laser melting (SLM), laminated object manufacturing (LOM) is also a topic for future research.AcknowledgementsThe research performed here was supported by funding from New York State Pollution Prevention Institute (NYSP2I) .References[1] Rock, S. J., and Wozny, M. J., 1992.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEnergy Flow (Bond Graph) Dynamic Modeling of Cartesian-Frame FFF 3-D Printer GantryMaharshi A. Sharma1 and Albert E. Patterson1,21Department of Mechanical Engineering, Texas A&M University, College Station, TX 77843 2Manufacturing and Mechanical Engineering Technology, Department of Engineering Technology and Industrial Distribution, Texas A&M University, College Station, TX, 77843maharshiarindom@tamu.edu; aepatterson5@tamu.eduAbstractEnergy flow (bond graph) modelling gives important information about the flow of energy to each component of a dynamic system and is especially useful for complex non-linear mechanical systems. This work presents a systematic development of a bond graph model of fused filament fabrication (FFF) 3D printer gantry. The model incorporates structural and belt stiffness, damping and input torque.

The model incorporates structural and belt stiffness, damping and input torque. The model was checked for correctness and causality using the 20-SIM software. The model was further validated using MATLAB-Simulink using parameters obtained for an example printer characterized in a lab environment. The bond graph model gives a unique view into modelling of the extruder carriage dynamics in FFF and can be applied to specific problems. It will also give interesting information on the controllability and system integration of the printer hardware.Keywords: FFF process; extrusion-based additive manufacturing; bond-graph model; dynamic modeling1. IntroductionOne of the common and useful ways to model mechanical systems is via the bond graph method, where exchanges of energy are modeled to simulate the behavior of the system. The method is discussed in depth in [1-4]. Bond graphs can also be used to generate algorithms to monitor operations in industries [5].

The method is discussed in depth in [1-4]. Bond graphs can also be used to generate algorithms to monitor operations in industries [5]. It provides a way of pictorial representation of the entire system consisting of multiple disciplines [6].Figure 1: Prusa frame printer system [16]. Adapted per the terms of a CC-BY license.One of the most common additive manufacturing processes is the fused filament fabrication (FFF) process, which uses a heated nozzle to selectively extrude thermoplastic feedstock to build parts up in layers. This is typically done using a cartesian-based 3-axis system. The basic machine has two major components moving relative to each686other, the extruder head, and the build plate, within the x-y-z coordinate system [7-10]. This current research focusses on Prusa-frame printers where the extruder carriage moves in X axis and Z axis direction and the build plate moves in Y axis direction.

This current research focusses on Prusa-frame printers where the extruder carriage moves in X axis and Z axis direction and the build plate moves in Y axis direction. Figure 1 shows the representative diagram basic Prusa-frame printer system and its axes of motion in coordinate system. This design allows the fastest printing speed of any of the cartesian-frame printers, is the easiest to use with thin layers (50-100μm) and has the lowest sensitivity to filament quality [11]. A clear understanding of the different types is essential in the correct modeling of the system to be applicable for modeling the extruder carriage motion.Bond graph presents an effort and flow representation of the system depicting the flow of energy from one component to another. The 0 in the bond graph are common effort (Force) and 1 are the common flow (velocity) junctions. 0 junction is generally used for velocity distribution and 1 junction is used for energy distribution.

0 junction is generally used for velocity distribution and 1 junction is used for energy distribution. The half arrow indicates the positive and negative directions of flow. I represent inertia; R represents resistance, in this study it represents damping; C represents compliance or 1/k. TF represents transformer, in this study it represents a ratio to be multiplied along the flow; and SE represents effort, for the purpose of this study it represents the external force applied.2. System Model Development 2.1 Printer ModelThe FFF system has three major components [7-10], the frame, the extruder carriage, and the build plate. For the purposes of the present study, the motion of the build plate is neglected and assumed to be static relative to the extruder head.

For the purposes of the present study, the motion of the build plate is neglected and assumed to be static relative to the extruder head. The plant for the rest of the printer can be divided into three major systems, each with a compliant connection; these are the (1) frame support structure (desk or workbench), (2) the printer frame, and (3) the extruder motion system.Figure 2: FFF machine plant configurationThe mount is connected to the system ground, while the energy input for the system flows into the extruder system. The energy for extruder carriage motions comes from motors mounted within the extruder motion system. These motors drive the mechanical components and induce the vibrations into the rest of the system. After further development of the system model shown in Figure 2, the spring-mass-damper model of the plant used in the extruder carriage motion can be derived (Figure 3) in x-z coordinates.

The definitions for each of the variables and parameters are shown in Table 1.Figure 3: Extruder carriage dynamic model687Table 1: Extruder carriage dynamic model parametersState VariableParameter𝒙𝟏 𝒙𝟐 𝒙𝟑 𝒙𝟒 𝜽𝟒 𝜽𝟓Position of printer mount (𝑚)Position of frame (𝑚)Position of extruder motion system (𝑚)Position of extruder carriage (𝑚)Angular position of Pulley 4 (𝑟𝑎𝑑𝑖𝑎𝑛𝑠)Angular position of Pulley 5 (𝑟𝑎𝑑𝑖𝑎𝑛𝑠)Parameter𝒎𝟏 Mass of printer mount (𝑘𝑔) 𝒎𝟐 Mass of printer frame (𝑘𝑔) 𝒎𝟑 Mass of extruder motion system frame (𝑘𝑔) 𝒎𝟒 Mass of Pulley 4 (𝑘𝑔) 𝒎𝟓 Mass of Pulley 5 (𝑘𝑔) 𝒎𝟔 Mass of extruder carriage (𝑘𝑔)𝒌𝟏 𝒌𝟐 𝒌𝟑 𝒌𝟒 𝒌𝟓 𝒌𝟔 𝜷𝟏 𝜷𝟐 𝜷𝟑 𝜷𝟒 𝜷𝟓 𝜷𝟔 𝒛(𝒕)Stiffness between ground and mount (𝑁/𝑚)Stiffness between mount and frame (𝑁/𝑚)Stiffness between frame and carriage motion system (𝑁/𝑚)Stiffness of belt section 1 (𝑁/𝑚)Stiffness of belt section 2 (𝑁/𝑚)Stiffness of belt section 3 (𝑁/𝑚)Energy dissipation between ground and mount (𝑁.

𝑠/𝑚)Energy dissipation between mount and frame (𝑁. 𝑠/𝑚)Energy dissipation between frame and motion (𝑁. 𝑠/𝑚)Energy dissipation in belt section 1 (𝑁. 𝑠/𝑚)Energy dissipation in belt section 2 (𝑁. 𝑠/𝑚)Energy dissipation in belt section 3 (𝑁. 𝑠/𝑚)Extruder carriage height (𝑚𝑚)The equations of motion for this dynamic system are derived using Newton-Euler method.

Bond-Graph ModelThe bond-graph model of the system is developed systematically to provide a better understanding of the sub-system and the flow of energy through the components. The printer system has been divided into system level blocks as shown in Figure 1 and into isolated extruder system, which when added gives the printer configuration.688(1)(2)(3)(4)(5)(6)The bond graph model in Figure 4 shows how the force applied at block 𝑚3 is transmitted throughout the system. The spring and damper system between blocks 𝑚1 and 𝑚2 and between blocks 𝑚2 and 𝑚3 acts on relative velocity between these blocks. The velocity is divided using the 0 junction while 1 junction represents the common velocity points connecting the springs and dampers of same relative velocity.

The velocity is divided using the 0 junction while 1 junction represents the common velocity points connecting the springs and dampers of same relative velocity. The 1 junction also connects the masses and is used to analyze the velocity of these masses.Figure 4: Printer block system (above) and bond graph model of the system (below)Figure 4, 5 and 6 shows a series of simplified mechanical systems that, when combined, constitute the mechanical system of the FFF extruder carriage (similarly to the method discussed in the Newton-Euler model derivation); note that 𝐷 is the generic connection force between the systems when they are connected during use. The most basic component is the simple drive belt system with two equally sized and equally massed pulleys and no slippage, shown in Figure 5a. The drive belt system is similar to the gantry system without the printer head. It is assumed that 𝑇1 > 𝑇2, which will be important in assigning directions in the bond graph model.

The drive belt system is similar to the gantry system without the printer head. It is assumed that 𝑇1 > 𝑇2, which will be important in assigning directions in the bond graph model. Figure 5b and 5c represent the bond graph model of the pulley system. Both these bond graphs represent the same system, giving the same results. The bond graph in Figure 5b is derived using efforts approach showing a step-by-step flow of energy or force throughout the system.

The bond graph in Figure 5b is derived using efforts approach showing a step-by-step flow of energy or force throughout the system. Figure 5c is derived using the flow approach depicting the division of velocities at various junctions in the system.689Figure 5: a) Drive belt system, b) bond graph of the system using efforts approach, c) bond graph model of the system using flow approach690Figure 6: a) Extruder system, b) bond graph of the system using efforts approach, c) bond graph model of the system using flow approachIf Figure 5a is modified to include a carriage, the result of the model will be Figure 6a representing the model of extruder carriage system. Figures 6b and 6c represent the bond graph model of the extruder carriage system.

Figures 6b and 6c represent the bond graph model of the extruder carriage system. The bond graph in Figure 6b is derived using efforts approach similar to figure 5b; Figure 6c is derived using the flow approach similar to Figure 5c.691Combining the bond graph of these sub-systems in Figure 5 and 6 provides the bond graph of the entire 3D printer (Figure 7). The dashed rectangular box in the bond graph in Figure 7 is similar to the bond graph shown in Figure 4. The flow method approach can be easily derived and is not shown here, although it is very similar to combining the bond graphs in Figures 4 and 6, while following the causality notation in the intersection shown in Figure 7. The torque applied into the system is through the pulley inside block 𝑚3, thus rotating the pulley which in turn moves the carriage. Applying Newton’s Law into the system, it can be observed that the block 𝑚3 will move in the opposite direction of the carriage.

Applying Newton’s Law into the system, it can be observed that the block 𝑚3 will move in the opposite direction of the carriage. Hence the flow and causality notation direction in the intersection between the extruder carriage system and the block system is clearly represented. This bond graph (Figure 7) can be used to measure both absolute and relative velocity of the extruder carriage system when it is in motion. The causality directions and bonds must be carefully considered and numbered in order to show the correct way to analyze or simulate the system using this method.Figure 7: 3D printer system (above) and the bond graph of the entire 3D printer system692Figure 8: Screenshot of the full system bond-graph model in 20-SIMOne of the features of bond graph models is the ability to directly produce the equations of motion for the system or its state space equations [6].

However, this exercise can be very impractical for large systems such as the one presented here, and the bond graph is better used directly in simulation of the system [1] using one of the many tools available for this purpose. The purpose of the developing this bond graph model is to give an additional view into the modeling of the extruder carriage dynamics in FFF and to provide an additional tool which can be applied (or modified as needed) to specific problems. All the models presented were carefully checked for correctness and proper causality using the 20-SIM® tool (University of Twente, Enschede, Netherlands), using simple inputs to ensure it behaved as expected. Figure 8 shows a screenshot of the bond-graph model in 20-SIM.3. System SimulationThe bond graph model of the sub-systems and the complete system have been shown in Figures 4-7.

Figure 8 shows a screenshot of the bond-graph model in 20-SIM.3. System SimulationThe bond graph model of the sub-systems and the complete system have been shown in Figures 4-7. While bond graphs can accommodate variable stiffness in the models, for the sake of this study, it was assumed that the stiffness and damping coefficient of the system were constant. The causality of the bond graphs was already verified using 20-SIM software in deriving the models. The equations of motion of the sub-systems and the complete systems shown were derived using Newton-Euler method to construct a MATLAB-Simulink model. The Simulink models for the bond graph systems are shown in the Appendix. The motion of the systems in the bond graph model were compared with the motion of the dynamic model in Simulink to validate the bond graphs. Note that arbitrary values of the parameters not related to this study were used to compare the results.

Note that arbitrary values of the parameters not related to this study were used to compare the results. Figure 9 shows the velocity of each block of the block system shown in Figure 4. Figure 10 represents the velocity of the pulleys and the carriage of the extruder system mentioned in the bond graph in Figure 6. The velocity plot of the complete 3D printer system is shown in Figure 11 and its corresponding bond graph is mentioned in Figure 7. In Figure 11 only the velocity of m3 block and extruder carriage is shown to verify the link between the sub systems. The plot in Figures 9-11 is obtained through bond graph using 20-SIM and the bottom plot is obtained through equation of motion of the system using Simulink.

The plot in Figures 9-11 is obtained through bond graph using 20-SIM and the bottom plot is obtained through equation of motion of the system using Simulink. It can be clearly seen that both these methods produced the same results for the block system, extruder system and the complete 3D printer system.693Figure 9: Velocity vs time plot of the block system obtained in 20-SIM (above) and Simulink (below)694Figure 10: Velocity vs time plot of the extruder system obtained in 20-SIM (above) and Simulink (below)695Figure 11: Velocity vs time plot of the 3D printer system obtained in 20-SIM (above) and Simulink (below)4. Conclusions and Future WorkThe Bond-graph model presented here provides a different insight into the system with interesting information of the hardware integration through its energy flow model. The bond-graph model developed represents the 3D printer in study accurately and can be used to track the extruder gantry motion.

The bond-graph model developed represents the 3D printer in study accurately and can be used to track the extruder gantry motion. The bond graph model developed can be used to represent a prusa-frame cartesian coordinate FFF 3D printer in complex industrial systems as well as to study the motion and flow within the system. The advantage of bond-graphs is that they can be extended to incorporate electro-mechanical systems and non-linear systems for a comprehensive study of dynamic systems. Bond-graph model can be further developed to incorporate non-linear parameters as well as motion in all 3 dimensions.AcknowledgementsThe authors thank Dr. James Allison for discussion and advice on formulation and dynamic modeling in this problem. While this publication represents a complete re-derivation and re-formulation of unpublished earlier work, the authors recognize Dr.

ERGY INPUTS TO ADDITIVE MANUFACTURING: DOES CAPACITY UTILIZATION MATTER?M. Baumers*, C. Tuck*, R. Wildman*, I. Ashcroft* and R. Hague*Additive Manufacturing Research Group, Wolfson School of Mechanical and Manufacturing Engineering, Loughborough University, Loughborough, LE11 3TU, UKAbstractThe available additive manufacturing (AM) platforms differ in terms of their operating principle, but also with respect to energy input usage. This study presents an overview of electricity consumption across several major AM technology variants, reporting specific energy consumption during the production of dedicated test parts (ranging from 61 to 4849 MJ per kg deposited). Applying a consistent methodology, energy consumption during single part builds is compared to the energy requirements of full build experiments with multiple parts (up to 240 units).

Applying a consistent methodology, energy consumption during single part builds is compared to the energy requirements of full build experiments with multiple parts (up to 240 units). It is shown empirically that the effect of capacity utilization on energy efficiency varies strongly across different platforms.IntroductionResearchers have argued that the world population‟s current ecological footprint far exceeds the planet‟s long-term capacity (Westkämper et al., 2000; Jovane et al., 2008). Energy generation and industrial activity contribute significantly to the overall emission of greenhouse gases (DECC, 2010), which are thought to be the key driver of global warming. A reduction of manufacturing energy consumption would thus be highly relevant for the limitation of overall greenhouse gas emissions. In this context, an understanding of the energy inputs consumed by the available manufacturing processes is critical. Foran et al.

In this context, an understanding of the energy inputs consumed by the available manufacturing processes is critical. Foran et al. (2005) have remarked that “if you can‟t measure, you can‟t manage”.The term „additive manufacturing‟ describes the use of a collection of technologies capable of joining materials to manufacture complex products in a single process step. Moreover, the technique allows the production of multiple components in a parallel manner (Ruffo et al., 2006) entirely without the need for tooling (Hague et al., 2004). The single-step nature of the additive processes affords an unprecedented level of transparency with respect to the energy inputs employed in the manufacture of complex end-use components. This advantage has received little attention in the literature and motivates the current paper.

This advantage has received little attention in the literature and motivates the current paper. Due to the parallel characteristics of additive processes, allowing the contemporaneous production of multiple parts in a build, it must be expected that the degree of capacity utilization during additive production impacts the total energy inputs. Hence, any corresponding summary metrics of energy consumption, such as the specific energy consumption per kg of material deposited, are also likely to be affected.the measurement of the electric energy inputs to additive processes. This approach is able to accommodate some of the differences exhibited by the various additive technology variants available in the marketplace.

This approach is able to accommodate some of the differences exhibited by the various additive technology variants available in the marketplace. Such variations include:  differences in build volume size, The current paper proposes a universally applicable methodology forthe ability to produce fully three-dimensional build volume packing configurations,a requirement for pre-heating and/or cool-down procedures.It is important to note that due to differences in terms of build material, layer thickness, mechanical properties and surface finish, the results reported in the current paper are not useful for direct comparisons of the energy efficiency (or environmental performance) of individual additive platforms.

The results reached in this paper do however offer a consistent and reliable absolute measure of energy consumption across a number of widely adopted additive manufacturing systems.The current paper concentrates on the spectrum of additive technology variants suited for the production of end-use parts and products, with an emphasis on metallic additive technologies. The following technology variants have been assessed experimentally for this research: selective laser melting (SLM), direct metal laser sintering (DMLS), electron beam melting (EBM), laser sintering (LS) and fused deposition modeling (FDM) of polymers. Table 1 summarizes system type, manufacturer reference, operating principle and nominal build volume size of the machines assessed in the performed experiments. Also, an indication of the layer thickness and build material selected for the performed power monitoring experiments is provided.SystemManufacturer referenceOperating principleNominal build vol.

In most cases, the literature provides an indication of specific energy consumption expressed in energy used (in MJ or kWh) per kg of material deposited. Table 2 provides an overview of the cited energy consumption results; where necessary, values were converted from kWh to MJ. Further, the table states if the data are derived from build experiments or estimated using secondary data. Regarding the work based on experimental results, it also reports whether the data were obtained from build configurations holding only a single test part or from builds with multiple parts. As the current paper will show, this is an important point.StudyTechnology variantEnergy consumption result MethodologyStereolithography74.52 – 148.97 MJ / kgLuo et al. (1999)LS107.39 – 144.32 MJ / kgEnergy consumption not empirically measuredFDM83.09 – 1247.04 MJ / kg3D Printing7.56 – 13.68 MJ per partMognol et al.

(2010a; 2010b) † - Calculated from data provided by Baumers et al. (2010)Table 2: Specific energy consumption results for additive processes in the literatureThe specific energy consumption results reported in the literature for the same additive technology variant can differ significantly, as noted by Telenko and Seepersad (2010), who suggest that differences in Z-height and density of the build experiments are responsible. This supports the assumption that the degree of capacity utilization is very likely to have a bearing on energy requirements. Thus, a consistent methodology in energy input measurement is needed for the analysis of process energy consumption.

Thus, a consistent methodology in energy input measurement is needed for the analysis of process energy consumption. This also prompts an important question about the efficient operation of additive manufacturing systems: to what extent does the degree of capacity utilization matter for each additive technology variant?MethodologyIf the parallel nature of additive technology is ignored in measurements of specific energy consumption, the usefulness of results may be impaired. To produce consistent results, such data should ideally be collected from experiments with a controlled degree of capacity utilization. The approach taken in the current paper rests on the assumption that for the efficient operation of additive manufacturing processes, it is necessary to produce parts in fully utilized build volumes, thereby operating the machinery at maximum capacity. This premise is confronted with the empirical data collected.

This premise is confronted with the empirical data collected. The underlying experimental strategy is to record the energy consumption during two specifically designed build experiments for each additive technology variant:In a first build experiment, the additive system is operated at full capacity. Where the technology‟s operating principle dictates that all parts are attached to a (removable) build plate or substrate, this is achieved by placing as many test parts as possible onthe available substrate area. Hence, in this case the available build capacity is exhausted in the X/Y plane. For technology variants allowing an unconstrained three-dimensional placement of parts in the build volume (for example LS), full capacity operation necessitates a three dimensional workspace configuration that uses up all available space in the X/Y/Z dimensions.The second experiment surveys the production of a single test part located in the center of the build volume floor plane.

This experiment provides information on the energy consumption characteristics if the available capacity is only minimally used, thereby allowing an analysis of the impact of capacity utilization on process energy efficiency.Due to additive manufacturing platforms normally being single-machine electricity driven systems, the measurement of electric energy inputs to build experiments is not complex. For the current research, process energy consumption was recorded using a digital power meter (Yokogawa CW240) attached to each system‟s AC power supply. Energy consumption was monitored throughout the entire build process, including any necessary process steps preceding and following the actual build activity. This includes, for example, process elements such as bed heating or vacuum drawing.

This includes, for example, process elements such as bed heating or vacuum drawing. With regards to the data gathered with the power meter, the focus lies on mean real power consumed per one-second measurement cycle (measured in W) and total cumulative electric energy consumed (measured in Wh). This cycle length is selected as the Yokogawa CW240 generates a full dataset of all available 137 measurement variables in this setting.The implementation of power monitoring experiments with consistent packing efficiency is based on the use of a standardized test part, shown in Figure 1. A reason for the „spider‟ shape is that it has a relatively large footprint in the X/Y dimensions, thereby limiting the achievable overall packing density and improving the economy and manageability of the experiments.

Due to the two-dimensional method of build volume packing found on some additive platforms, this is particularly effective for approaches that require every part to be attached to the build substrate.Figure 1: A standardized test partA further consideration in the design of the test part is that it should not require auxiliary structures for the support of overhanging areas or the dissipation of heat. In the build experiments on the MTT SLM250 and the Stratasys FDM 400 mc, support structures were generated to connect the parts to the build plate. This configuration was chosen idiosyncratically for the MTT SLM250 and the build could have been performed directly on the substrate, therefore the energy expended for the supports is not factored into the energy consumption results. On the FDM 400 mc, parts are normally connected to the removablesubstrate using a secondary support material.

On the FDM 400 mc, parts are normally connected to the removablesubstrate using a secondary support material. For this reason, the energy consumption results include the generation of support structures on this platform. For all analyzed platforms, the energy consumption of ancillary equipment, such as optics chillers and post processing equipment (e.g. used to cut the parts off the build plates) is not considered.In contrast to the simplicity of the power monitoring setup, the creation of a standardized level of full build capacity utilization across different additive platforms proved non-trivial. The idea pursued in this research was to apply a custom-developed build volume packing algorithm, producing densely packed, albeit sub-optimal, build configurations for use in the full build experiments. The resulting implementation is capable of generating build configurations both in full 3D mode as well as in a 2D mode constrained to the substrate area.

The resulting implementation is capable of generating build configurations both in full 3D mode as well as in a 2D mode constrained to the substrate area. A simple algorithm is used to insert and translate voxel representations of the test part in the build volumes in order to produce a densely packed configuration. In the full 3D packing mode, used only for the full build on EOSINT P 390, the algorithm was also allowed to flip test parts upside-down.the surveyed metallic additive systems exhibit an approximately square horizontal cross-section with a similar side length ranging from 200 mm (Arcam A1) to 250 mm (EOSINT M 270 and both SLM systems). However, the presence of rounded build volume corners and holes in the substrates made the algorithm- generated build configurations unacceptable.

However, the presence of rounded build volume corners and holes in the substrates made the algorithm- generated build configurations unacceptable. Thus for the metal platforms, it was decided to use human operator packed builds in the full build experiments, resulting in a total number of five (Arcam A1) and six (EOSINT M 270 and both SLM systems) test parts included, as shown in Figure 2. While all metallic additive platforms feature atmosphere generation and (in some cases) pre-heating routines, only the EBM platform runs a controlled cool-down procedure after each build.The build volumes ofFigure 2: Full build configuration for SLM and DMLS (left) and EBM (right)Both full build experiments on polymeric systems were successfully configured with the build volume packing algorithm. As the EOSINT P 390 features a very tall build envelope (nominal Z-height: 620 mm), it was decided to limit the build experiment to a 50 mm horizontal „slice‟ of the available build space.

As the EOSINT P 390 features a very tall build envelope (nominal Z-height: 620 mm), it was decided to limit the build experiment to a 50 mm horizontal „slice‟ of the available build space. The resulting packing configuration (holding 20 test parts) is shown in Figure 3. The full build energy consumption result for this system was obtained by extrapolating the energy consumption data measured during the 50 mm build phase to the full 600 mm of available build height. No adjustments were made to the energy consumption increments added during system warm up (160 min) and cool down (914 min).Figure 3: 3D build configuration for EOSINT P 390, 50 mm Z-heightFigure 4: Build configuration for FDM, nominal build volume area: 406 × 355 mmFor the build experiment on the FDM 400 mc, the packing algorithm was used in the two-dimensional mode, this resulted in the insertion of 16 parts (Figure 4).

A further characteristic of the FDM 400 mc affects the energy consumption results: as the interchange of substrates is carried out in a fully heated machine and the machine is designed for continuous operation, zero warm-up and cool-down time and energy consumption are assumed.Results and DiscussionThe current research pursues two goals: firstly, it aims to contribute reliable summary data of additive process energy consumption in an inter-platform study. As the density of the parts in the build volume is expected to play an important role, an emphasis is placed on keeping the degree of capacity utilization constant across platforms (although no attempt is made to quantify this). This additional clarity is needed, as the existing literature reports a wide range of energy consumption levels, even for technologically closely related platforms.

This additional clarity is needed, as the existing literature reports a wide range of energy consumption levels, even for technologically closely related platforms. Secondly, the application of the above described methodology allows an analysis of the impact of variation in the degree of capacity utilization - comparing production in a state of incomplete utilization (one test part per build volume) to production at exhausted capacity (fully occupied build volume).In an effort to understand the determinants of additive process energy consumption, it is instructive to compare mean real power consumption throughout the performed full build experiments. Figure 5 shows that the variation of mean power consumption observed for theassessed platforms is not very large, ranging from 1.09 kW to 3.33 kW, with a standard deviation of 0.76 kW. This is in contrast to the extensive differences in specific energy consumption reported in the literature (52.20 to 1247.04 MJ per kg).

This is in contrast to the extensive differences in specific energy consumption reported in the literature (52.20 to 1247.04 MJ per kg). It should be noted that a contributing factor to the small mean power consumption exhibited by the MTT SLM250 (1.09 kW) is that the external optics chiller draws power from an external source (~ 0.6 kW) and does not appear in the data.Figure 5: Mean real power consumption during the build experimentsFigure 6 reports the specific energy consumption per kg of material deposited during the build experiments for both the single part (dark column) and full build (light column) experiments. For all surveyed systems, as expected, full capacity utilization results in a lower specific energy consumption.

For all surveyed systems, as expected, full capacity utilization results in a lower specific energy consumption. However, the size of this saving varies heavily from platform to platform.Figure 6: Single part and full build energy consumption per kg depositedIt appears that for the energy efficient operation of some additive systems, capacity utilization is critical. The two polymer-processing systems form the extremes in this comparison. While the specific energy consumption observed on the FDM 400 mc appears relatively insensitive to the switch to full capacity utilization (-3.17 %), the change to full utilization produces a huge specific energy saving on the EOSINT P 390 (-97.79 %).Contributing factors to the disproportionately large variation are likely to be the extensive energy investments in machine warm-up and cool-down. As Table 3 shows, these energy investments are large during both build experiments on the EOSINT P 390.

As Table 3 shows, these energy investments are large during both build experiments on the EOSINT P 390. The much greater number of parts generated in a full build configuration (240) enables the listed energy consumption to be allocated to a far greater part mass, thereby producing a favorable process energy consumption result (107 MJ per kg).Single part buildFull build experimentBuild phaseEnergy consumptionDurationEnergy consumptionDurationWarm-up31.55 MJ125 min37.96 MJ160 minCool-down66.94 MJ600 min97.33 MJ914 minTable 3: EOSINT P 390 warm-up and cool-down energy consumption and timeContrary to this, the FDM 400 mc is assumed to operate continuously. Therefore, extra energy consumption increments for build volume warm-up and cool-down are unnecessary. The substrate carrying the produced parts is removed from the machine at operating temperature and replaced by an empty substrate for the next build.

The substrate carrying the produced parts is removed from the machine at operating temperature and replaced by an empty substrate for the next build. This, of course, ignores periods in which the machine may in practice sit idle due to other reasons.The metallic platforms exhibit a smaller variation in specific energy consumption. Operating at full capacity, the Arcam A1 exhibits a far greater energy saving (-65.54 %), than the MTT SLM250 (-21.70 %). Again, the reason for this is likely to be the significant energy expenditure for build volume pre-heating and cool-down procedures. While consuming markedly more energy per kg of material deposited than the other metallic systems, the EOSINT M 270 and Concept Laser M3 Linear, show a similar variation in specific energy consumption (-28.91 % and -28.06 %).

A possible reason for the comparatively high specific energy consumption levels observed during both experiments is the small layer thickness setting used on these platforms (20 µm and 30 µm).The presented results show that the change to full capacity operation results in a reduction of the energy consumed per kg of material deposited for all assessed operating principles and build materials. This gives support to the proposition that the energy consumption data derived from full build experiments are reflective of technically efficient machine operation. Full capacity production should therefore serve as the yardstick in the evaluation of process energy efficiency. Considering that the differences in mean real power consumption do not appear large across the surveyed platforms, the prime candidate for the determination of specific energy consumption of additive processes appears to be machine productivity.

A direct relationship between process productivity and specific energy consumption is proposed by Luo et al. (1999). The specific energy consumption rate in MJ per kg deposited is denoted by SEC, process productivity is symbolized by PP and PR denotes the mean real power consumption (power rate), such that:(1)A scatter plot (Figure 7) of the full build results against the process productivity measure illustrates the negative relationship between machine speed and energy efficiency.The results for the single part build experiments were not included in Figure 7 as they are not seen to be reflective of efficient technology utilization.Figure 7: Specific energy consumption versus machine productivityConclusionsThe heterogeneity in the approaches taken in the literature towards the measurement or estimation of the specific energy consumption of additive machines highlights the need for the development of a more consistent methodology.

The technique presented in the current paper allows the calculation of specific energy consumption metrics for different platforms and systems, irrespective of the fundamental differences in operating principle and build materials. Even though all power monitoring experiments were based on the same test part geometry and some of the resulting parts look and feel quite similar, it is likely that the applications of these processes vary due to differences in material, mechanical properties, and surface quality. Therefore, these results should not be interpreted as statements on the relative environmental performance of the assessed additive technology variants.However, the current research has conclusively demonstrated that the realized degree of capacity utilization has an impact on process energy efficiency.

In the performed experiments, this impact on energy savings in terms of specific energy consumption per kg of material deposited ranged from small (-3.17 % for FDM) to extreme (-97.79 % for LS). Considering the LS and EBM processes, which include extensive energy expenditure for atmosphere generation, warm-up and cool-down, full capacity utilization will result in far greater energy efficiency compared to a single part mode of production. This may pose problems in the estimation of process energy consumption for additively produced parts if the composition of the production build is unknown. Contrasting this, the results indicate that the FDM process (where system warm-up and cool-down do not enter the energy consumption metric at all) does not benefit significantly from full capacity utilization. It appears that FDM can be applied in a serial fashion generating output part-by-part without incurring a significant energy efficiency loss.

hanced controlling of the SLS Process during a build.Alexandre L. Papadatos, Kenneth Stanley Jr., Said Ahzi and Frank W. PaulCenterfor Advanced Manufacturing Department of Mechanical Engineering, Clemson University, Clemson, SC 29634AbstractCurrent commercial Rapid Prototyping(RP) systems like Stereolithography (3D Systems Corporation) use galvanometers from General Scanning Inc. (GSI) for thepositi()ning of the laser beam. The GSI scanners are delivered as a ''black box". Operating Consoles which are usually Personal Computers ·(PC) have ·very few feedback· from the performance of the GSI scanners. Therefore, the PC spends 900;0 or more of its timewaitingJor the GSI scanners to be over <with the building of the current layer before sending the information regarding the next layer. Also, very little process controLcan be performed during the scanning of a layer using the GSI scanners.

Also, very little process controLcan be performed during the scanning of a layer using the GSI scanners. This kind of setup prevents any dynamic controlling of the process·thatcouldprevent building errors like burning, warping etc.and Selective Laser Sintering (DTM Corporation)At Clel11Son University, our team hasaevelopedbothhardwareand software components that allows a dynamic con.trol of the building process. New features like scanning one vector with laser power as a function of position and/ or time are now possible. Both hardware and software issues will be presented.BackgroundTo create objects using the Selective Laser Sintering process, an X Y optical scanner using CAD solid model data,guides or directs the laser beam to target points requiring sintering. TheSLS system employs the General Scanning Incorporated (GSI) optical scanner, the DE series 2000 scanner controller and two G325 DT galvanometers with mirrors (1,3).

TheSLS system employs the General Scanning Incorporated (GSI) optical scanner, the DE series 2000 scanner controller and two G325 DT galvanometers with mirrors (1,3). The objective of Stanley's work (1) has been to identify the technological limitations of the GSIoptical scanner, as related to the scanning requirements of SLS,and present a new optical scanner design based on digital signal (DSP) control theory. One important accomplishment of the research ..done by Stanley is the use of information derived ~om the closed loop position control. of the galvanometers to trigger the laser onloff function. This functionality has been enhanced and .constitutes one of the main components for the dynamic optical scanning controller.The GSI system has always been used as a ''black box", and a very small amount of data is returned to the operating console.

We will describe how the build of an object is managed: the software that supervises the SLS process (the main software) downloads the informations required to build one single layer into the GSI system. When the downloading is over, the main software orders the GSI system to go ahead and build (scan) the layer. The main software then waits for the GSI system to finish. When over, the GSI system sends an "OK" command back to the main software. Then the main software becomes aware of the factthafscanning operationsareovefforthatlayer. It then spreads a new layer of powder and downloads to the GSI •. system the informations required for the processing of the next layer.In the GSI system, the laser controlling software that processes the scanning of a layer cannot be willingly modified. It isa property of General Scanning Inc., and it is engraved in a memory chip.

It isa property of General Scanning Inc., and it is engraved in a memory chip. However, the new controller developed by Stanley (1), uses a47Texas Instrument digital signal processor board (4), which is capable of implementing bo~ digital control of the scanning system and supervisory control of the SLS process. A mam software is still required in order to read the commands for the layers, to perform. tI:e manao-ement of the controller operation and as a user interface with the system. This IS mainl~ because the. Texas Instru.ment board cannot access devices such as the hard drive, the communication ports and the monitor. Therefore, an interface between the board and these devices is needed, and this is the primary function of the main program in the new controller system: the dynamic scanning controller. The main difference with the GSI system is that.

The main difference with the GSI system is that. the software controlling the laser position, speed and intensity can now be modified to meet our needs.Hardware Designfor the· Dynamic Optical Scanning ControllerThe optical scanner developed by Stanley (1) utilizes the actuators •.and sensors of the GSlscanning systemrbut provides anew controller based on digital signal processing (DSP).With theDSP based controller, control of the laser tracking is accomplished with software and control algorithms instead of the hardware controller from the GSI system. This technique provides thettser with a flexible control environment for laser tracking and also allows control of laser on/offand laser intensity signals based on the galvanometer's position.

This technique provides thettser with a flexible control environment for laser tracking and also allows control of laser on/offand laser intensity signals based on the galvanometer's position. The sensors and actuators have been retained by Stanley in hisDSP based system because they were the best technology available ·to move the mirrors with the required performances for rapid prototyping optical scanning.The controller of theDSP optical scanner uses a C30 digital signal processor system from Spectrum Signal Processing, Inc. of Westborough, Massachusetts (3). The C30 system board is designed to coexist in an IBM compatible PC. Data transfer between the DSP and PC is accomplished through the ISA-ATbus. The C30 system board has 64K of 32 bit RAM memory that is accessible by both the PC and DSP. These 64K of memory will be split into two memory banks of 32K each, in order to speed up the process. This feature \lvill be discussed later in this chapter.

These 64K of memory will be split into two memory banks of 32K each, in order to speed up the process. This feature \lvill be discussed later in this chapter. More detailed information on the board itself can be found in (1,3,4).Since the DPS system uses the same sensors and actuators as theGSI system, an interface board was required. It has been successfully designed and developed by Stanley (1). The interface board hooks on the GSI system and allows the control of the SLS process either with the GSI system or the newer DSP system. All the signal of the DSP / GSI interface board are designed on one circuit board that is sized appropriately to fit in the EURO Card Frame of the DE 2000 series controller from General Scanning.

All the signal of the DSP / GSI interface board are designed on one circuit board that is sized appropriately to fit in the EURO Card Frame of the DE 2000 series controller from General Scanning. A schematic for the DSP/ GSI interface board and further technical information can be found in Appendix A of (1).48Software Design for the Dynamic Optical Scanning ControllerTerala (2) and Stanley (1) initially developed the control program developed for the CLEMSON sts project, using the GSI controller. The new controller program incorporates some of the control software developed by Terala and Stanley.

The new controller program incorporates some of the control software developed by Terala and Stanley. Our contribution consists of further testing the software, correcting some problems due to communications problem between the DSP and the PC, and also enhancing the capability of the DSP system by featuring the variation of the laser intensity while scanning a vector.The software developed for this research, uses scan commands with the same format as the GSI controller, in order to prOVide compatibility with previously generated scan files. One new command is added, called the Laser Setting. The syntax is LSxxxxx where "xxxxx" is an integer number from 0 to 65535. Each vector can now be scanned with a particular laser intensity pattern. While being scanned, each vector is divided into ten zones, each equal in length. One laser intensity can be specified for each of these ten zones. Eight different levels define the laser intensity.

While being scanned, each vector is divided into ten zones, each equal in length. One laser intensity can be specified for each of these ten zones. Eight different levels define the laser intensity. The first level (level "0") represents 2% of the maximum laser intensity (Le. 1 Watt for a 50W C02 laser). The second level (level "I') represents 16% of the maximum laser intensity, and so on. At the end, level "7" represents the maximum intensity that can be provided by the laser apparatus. This configuration was chosen in order to be able to code the pattern in an octal form. The octal format has been retained since the decimal format would require the use of three LS commands for each scan vector, making a total of five commands required to build a single vector. Choosing a smaller base would not provide such a wide range of power levers.

Choosing a smaller base would not provide such a wide range of power levers. The use of the octal format requires only two LS commands to describe the way a vector should be scanned, bringing the total to four commands for each scan vector. Figure 1 shows five vectors, scanned each with an independent scanning pattern. We have scanned these vectors on paper in order to keep a visual track of the change in laser intensity. The five vectors are horizontal and all are equal in length.. On the top of each vector we have added the laser intensity level for each of the laser zones. The zones are shown at the top.777777777012WZ?345677765432100024677642~"'*707070707Figure 1. Horizontal vectors with different scanning laser pattern.In the encoding of the scanning pattern describing these five vectors, two LS commands are used for each of the vectors.

Horizontal vectors with different scanning laser pattern.In the encoding of the scanning pattern describing these five vectors, two LS commands are used for each of the vectors. The numbers following the two LS COmmands describe the pattern. For instance, the two first LS commands found in the script file are "LS16383" and "LS65535". They correspond to the first vector, the one on the top of Figure 1. In order to derive these numbers, we first choose the desired pattern. For the first vector the desired pattern was vector with maximum and uniform laser intensity. The maximum intensity is level 7; therefore all the ten zones are assigned level 7. In an octal form we have49770077777777777 describing the desired pattern. Ifwe subdivide this octal number in two integer numbers, of 16 bit each, we get in binaryform 11111111111111 for the first integer number (i.e. 16383 in decimal) and 1111111111111111 for the second one (i.e.

16383 in decimal) and 1111111111111111 for the second one (i.e. 65535 in decimal).the desiredpattem is 0123456min ..octal form or 000001010011100101110111111111.inbinaryfo.rm.We divide this binary number into two integer 16 bit numbers: 00000101001110 (Le. 334) and 101110111111111 (i.e. 24063). Thus, the LS commands·for the second vector are "LS334" and 'LS 24063".Eo.rthe second vector,Computer ImplementationThe control software of the. DSPcontroller has two parts due to the design of the system (1). The DSP processor requires the support of a hoslcompllter for data storage, userin.terfaceandcommunication with the SLS machine. Stanley named the program for thehostPC "PCBUILD.C" and for theDSP system "DSPBUILD.C". ThePCprogramhas been written using Borland™ C++ v45. When the program<starts, a xnenu is displayed on the screen. The user can perform from there several actions. In order to start building a part,the use:.

When the program<starts, a xnenu is displayed on the screen. The user can perform from there several actions. In order to start building a part,the use:. needs to initialize the DSP system by selecting the first option of the menu. Detailed instructions on how to use the PC program can be fOllnd in (6). The DSP program has been written inC using a text editor and the special compiler provided by Texas Instrument. Most of the program is similar·to·the version written by Stanley. However, some communication problems between the PC and the DSP had risen as the entire system was used and tested. In order to prevent comrrnmication problems, we have enhanced the protocol used for the DSP and the PC, by adding stages during the communication, where the DSP and PC· processors have acknowledge of·each other's status: The DSP waits on the PC to acknowledge its· status, before .continlling. Then the PC waits its turn on the DSP to acknowledge its status.

Then the PC waits its turn on the DSP to acknowledge its status. Due to <this added checkings,the time for initialization and data transfer has been slightly increased compared to Stanley's version. However, communication problems that caused the system to freeze, have virtually disappeared.Another enhancement over Stanley's version of the DSP program, is the addition of the laser intensity monitoring. Since theDSP system supervises the scann.irl.g of each vector, the system is also capable on setting the appropriate laser intensity during the scan of a vector. The actions corresponding. to theLS commands have been implemented;. The system monitors the scan of a vector and sets the laser intensity accordingly to the value specified after the. two· LS commands. In .order for the LS commands to be considered, the commands "LPO" and "LRO" need to be previously specified.

two· LS commands. In .order for the LS commands to be considered, the commands "LPO" and "LRO" need to be previously specified. This restriction was chosen to prevent confusion of the system by opposite commands: LP and·LR commands specify scanning of vectors with uniform laser intensity, while LS commands specify vectors with variable laser intensity.An optimization concerning the time required for data transfer from the PC to the DSP has been initially designed by Stanley. The 64 kilobytes of memory that the DSP is equipped in order to stOre data have been divided into two banks of 32 kilobytes each. The reason that led us to this decision is that the informations for one layer are seldom larger than 32 kilobytes. This division of the DSP memory allows us to fill one memory bank with data from one layer, while we process the data of the previous layer stored in the other memory bank.

This division of the DSP memory allows us to fill one memory bank with data from one layer, while we process the data of the previous layer stored in the other memory bank. Therefore, in the ideal context where the time to process a layer is equa.l to the time to transfer the data into theDSP memory, the data transferring time is reduced by half. Practically, the laser is always scanning faster the layers than with the GSI system, since most of the informations for the next layer have already been transferred in the available memory bank when the laser ends a layer.Dynamic Controller ResultsIn order to test the ability of the new controller, we took an existing scan file and added manually LS commands for each of the vectors. We then ?roceed~d to the build using the DSP controller. Figure 2 displays the result of the dynarruc scannmg of one lay~r of a dogbone saIl1ple. We used paper as a Il1ediUIll in order to visually track the dynarruc scanning.b<F<Figure 2.

Figure 2 displays the result of the dynarruc scannmg of one lay~r of a dogbone saIl1ple. We used paper as a Il1ediUIll in order to visually track the dynarruc scanning.b<F<Figure 2. Dynamic scanningIn Figure 2, we can clearly see where the laser intensity changes. The bottom "dogbone" used a constant laser intensity to proove that dynamic scanning can behave similarly as when using the GSI controller. The three dogbones have been scanned using horizontal vectors. The dynamic controller takes each of these vectors and divide them into ten equal-in-Iength zones. The largest the vector, the largest the zones will be and th~ worst the resolution will get since we. can only specify one laser setting per zone. There. are two ways to remedy to this problem.:1) Increase the number of zones per vector. This would result in a direct increase of the length of the script file ~ince more LS commands will. be necessary.

This would result in a direct increase of the length of the script file ~ince more LS commands will. be necessary. A major update of the software is required in this case, including code programming.2) Scan the part in a different way. For instance, the use of vertical vectors instead of horizontal, will result in shorter vectors. Therefore, the zones will be shorter.and the resolution will also be better. This will lead to an increase of thelength of the scan and scriptfiles, since more vectors will be generated. In this latter. case however,noupdate.ofthe software is rEilquired.51The Building ScriptAt Clemson University, a program simulating the SLS process has been developed and implemented (5,7). Once a simulation has been performed, we are in possession of data representing the powderbed after the build thatcan be displayed on a computer screen as a visual representation of the powder bed.

Once a simulation has been performed, we are in possession of data representing the powderbed after the build thatcan be displayed on a computer screen as a visual representation of the powder bed. These visualrepresentations already provide the SLS user with important informations on how the build will actually look like. Among these informations are burning (thermal degradation of the material), bonus z and low sintering (5). However, without a tool that would modify dynamically the building parameters, the user can do very little in order to prevent these unwanted phenomena.The dynamic controller allows supervising the entire SLS process. The dynamic controller also allows varying the laser intensity settings while scanning a vector. Each vector can therefore have its. individual laser power setting depending on the geometry.

The dynamic controller also allows varying the laser intensity settings while scanning a vector. Each vector can therefore have its. individual laser power setting depending on the geometry. However, due to the amount of vectors required to build an entire geometry, specifying manually the laser setting for each vector becomes an impossible task. This difficulty is the main reason that led us towards the development of a new tool. We will present the to analysis software that uses the results from the simulation software in order automaticallyl generate a script file (an improved scan file) which will use the new features of the dynamic controller.Automatic Generation of a Scan FileFor a machine to perform a task.autornatically, ifis necessary to give the machine certain guidelines on how to perform in different situations it will encounter.

The set of the guidelines is what we will be calling the "analysis rules" and depending on them, the machine will perform accordingly in the analysis of the simulation results. The analysis rule we have adopted as a testbed for our system is based on the maximum temperature reached in the powder bed. We have been referring to the scan file used by the simulator and the GSI scanners as the "scan file". We will be referring to the new scan file that is generated by the analysis software as the "script file", since it also includes informations regarding the improvement of the build. The scan file and the script file are describing the same geometry. However, the scan file uses two commands to define a vector, whereas the script file describes a vector using four commands; the two additional commands describe the laser power settings for a given vector.

The result is that the script file is close to double the size of the scan file.For each scan vector describing the geometry, the following procedure is followed:1) We initialize all the vector zones with the laser setting specified in the input file. For the sake of simplicity, we have assumed that the laser setting is higher than needed. Therefore, the laser setting can only be decreased and the phenomenon of low sintering is not taken into consideration in the analysis rule presented here. 2) We make sure to identify the vector zone we are currently in, depending on the progression of the vector scanning. As we saw, there are ten vector zones, which are equal in length for each vector and seven laser power settings. 3) We call the subroutine AnalyzeO which can be found in the module models.c. This subroutine will decrease the laser setting in each vector zone according to the information from the simulation data on this region of the powder bed.

This subroutine will decrease the laser setting in each vector zone according to the information from the simulation data on this region of the powder bed. 521) We initialize all the vector zones with the laser setting specified in the input file. For the sake of simplicity, we have assumed that the laser setting is higher than needed. Therefore, the laser setting can only be decreased and the phenomenon of low sintering is not taken into consideration in the analysis rule presented here. 2) We make sure to identify the vector zone we are currently in, depending on the progression of the vector scanning. As we saw, there are ten vector zones, which are equal in length for each vector and seven laser power settings. 3) We call the subroutine AnalyzeO which can be found in the module models.c. This subroutine will decrease the laser setting in each vector zone according to the information from the simulation data on this region of the powder bed.

This subroutine will decrease the laser setting in each vector zone according to the information from the simulation data on this region of the powder bed. 524) Once a vector has been analyzed, we derive the two attributes that will be specified byth~ two "LS" commands. 5) Unless we are dealing with a scan vector, we copy the same command we read in the scan file into the script file. However, if we are dealing with a scan vector, we first write the two "LS" commands for the scan vector in the script file and then we copy the "NX" and "NY" command. This is in order for the dynamic controller to know a priori what laser setting is required to scan the given scan vector. Performance and LimitsThe analysis rule we have used for testing the system is based on the maximum temperature reached in the powder bed. The script file that is generated based on this analysis will only have an effect on the burnil1.g. This is mainly due to the simplicity of the analysis that we have used.

The script file that is generated based on this analysis will only have an effect on the burnil1.g. This is mainly due to the simplicity of the analysis that we have used. In order to take into consideration other problems, a more extended analysis is required, and more complex analysis rules need to be implemented. With a better-detailed analysis, the system including the simulation software, the analysis software and the dynamic controller will provide better quality of the parts made using the SLS process. However, the goal of the research presented in this thesis has been to prove the feasibility of such a system, and this is the reason we have used a simple analysis as a testbed. Figure 3 shows the part built 'With the GSI controller using a constant laser intensity setting.

Figure 3 shows the part built 'With the GSI controller using a constant laser intensity setting. Figure 4 shows the same part built using the dynamic controller in conjunctionwith the simulation and analysis software.VVhen using the GSI controller, the laser delivers too much energy into the powder bed as the length of the vector decreases at the edges. We c~ see that the burning at the edges is eliminated when using the dynamic controller. The burning was predicted by the simulation software, and the analysis software created a script.file in such a way that the laser intensity is decreased, as the vectors become shorter.Figure 3. Part built using the GSI controller.53Figure 4. Part built using the DSP controller.ConclUSionThe DSP system developed at OenlSonUnivetsity has been improvedintwo major ways.

Part built using the GSI controller.53Figure 4. Part built using the DSP controller.ConclUSionThe DSP system developed at OenlSonUnivetsity has been improvedintwo major ways. Some problems cause~by faulty cornmunicationbetween the DSP and PC programs have been corrected, and c~ntr0lling;thelaserintensitydefined by a new c.ornmand has been i1.TIplemented. One single vectof.can.bescannedusingits ownlaser intensil[pattern, independently of previous· or next configurati.ons. This feature allows t.o reduce the intensity when burning occurs specially at the edges, or to increase theinte11Sity if the sintering is not enougl:l.HQ~ever,~heJ:l()Jrl.y cOll$iders thf;)uumber.of vectors describing a three dimensional CAD ~~j~.qt'i<imanUal~tillg; ofithe~9an#le>~ecomesaniJ.npossible task.

19914. Spectrum Signal Processing Incorporated,"TMS320-C30 System BoardTechnical Rejerence Martual." Spectrum Signal, Westborough, MA 19925. Papadatos A. L., Ahzi, S., Deckard, C.R, Paul, F. W."On Dimensional stabilities: Modeling of the Bonus-Z During the SIS Process ", SFF Symp. Proc., 8,pp. 709-716, 19976. Davis A.L.,"User's Manual', Clemson University Mechanical EngineeringDepartment,. SLSLaboratory, version ·1.01, June 19987. Papadatos A.L.,"Computer Simulation and Dynamic Control of the SIS Process", Clel11son .University Mechanical Engineering Department, SLS Laboratory, August 19985556.

HANCING SIMULATIONS OF ADDITIVE MANUFACTURING PROCESSES USING SPATIOTEMPORAL MULTISCALINGD. Pal*Ϯ, N. Patil*Ϯ, C. Teng*Ϯ, K. Zeng*Ϯ, S.XuϮ, T. SubletteϮ and B. Stucker*ϮDepartment of Industrial Engineering, University of Louisville, Louisville, KY 40292. Ϯ3DSIM LLC, Louisville, KY 40202.AbstractThe three integral components which are common to all additive manufacturing machines are energy source(s), material(s) and geometrical representation(s). The interaction of these components lead to coupled multiscale thermomechanical phenomena. The overall response due to dynamic localized boundary conditions leads to travelling thermomechanical waves with their origin at the point/line or plane of energy input and finite boundaries located at the extremities of the build chamber of the machine. At these finite boundaries, three phenomena occur, namely reflection, absorption and transmission of the thermomechanical travelling wave.

At these finite boundaries, three phenomena occur, namely reflection, absorption and transmission of the thermomechanical travelling wave. Based on the boundary conditions, any combinations of the mentioned finite boundary phenomena are activated leading to a finite mismatch between the boundary condition and ordinate of the travelling wave at the finite boundary abscissae. This finite mismatch leads to another thermomechanical travelling wave which travels from the finite boundary towards the location of energy input and the process of generation of thermomechanical waves continues till a standing steady state thermomechanical wave converges at each snapshot in time. An important aspect of these forward and backward moving travelling waves is their spatiotemporal thermomechanical curvature as a function of space which leads to spatiotemporal multiscale modeling of these processes.

Spatial and temporal examples with multi-order and multi-rank updates of thermal eigensolutions with inhomogeneity in the –z direction and thermal stiffness as a function of time stepping were thus studied to understand their effects on efficient Finite Element simulations of metal additive manufacturing processes.Introduction and BackgroundAll Selective Laser Melting (SLM) process models are computationally expensive to solve using traditional FEM methods, as they require fine-scale resolution across a part volume that is many orders of magnitude larger than the fine-scale lengths. This has led to efforts to solve the problem using asymptotic behaviors, modal space methodologies or beam theories. The intent behind this new EigenSolver strategy is to solve for dominant modes in the macroscopic domain and local high frequency modes in the microscopic domain. These can be accurately applied at certain locations (e.g.

These can be accurately applied at certain locations (e.g. 5 layers) below the melt pool.Beam and plate theories are meant to take benefit of asymptotic theories involved in dimensionally reducible structures. Beam theories for complicated geometries such as aircraft wings and ship hulls and for material variations across the cross-section and length of a component have been an area of intense research. An analytical method, namely the Variational Asymptotic Method (VAM) [1], has been used extensively to derive these beam theories. One of the limitations of this methodology lies in the difficulty of deriving beam theories for very complicated beams. To overcome this problem we have developed a new EigenSolver beam theory derivationfor use with FEA. This novel method can consider any complicated shape and cross-sectional variation and derives beam or plate theories involved in it. Further it has applications in problems which do not fall in the categories of beams or plates (e.g.

Further it has applications in problems which do not fall in the categories of beams or plates (e.g. prismatic bodies or any general structural configuration). The importance of this methodology lies in calculating eigenvectors of the cross- section or group of nodes in FEM matrices (such as the stiffness matrix). It is well known that eigenvalues and eigenvectors are computationally expensive to calculate. To overcome this difficulty, any set of orthogonal vectors can be employed for this purpose and their coupling is calculated to derive simplified simultaneous equations. Various material and geometrical nonlinearities can be incorporated in this beam theory with the help of our other existing solver tools.This methodology can be further generalized to work with groups of nodes in banded FEM meshes for large size problems. These can then be solved on desktop computers.

These can then be solved on desktop computers. In summary this methodology is capable of:Representing complicated geometry and material inhomogeneity in beams.   Generalization to be used for any problem with some hidden unknown asymptotics.  Solving large size FEM problems (for example 10 million Degrees of Freedom or much larger in some specific cases) on desktop computers.Inexpensive computations leading to global-local coupled response.Simplified and fast assembly of FEM matrices.In addition, an analytical method using a multi-dimensional, multi-resolution wavelet based eigentheory is in development. This methodology can solve both linear and non-linear problems provided a space transformation between the cross-sections is available.

This methodology can solve both linear and non-linear problems provided a space transformation between the cross-sections is available. The theory has been extended to account for inhomogenous material distributions which are very common among powder based additive manufacturing technologies where solid and powder continuously interact in a non-linear mode with the applied thermal and resulting deformation based boundary conditions.A schematic diagram which represents the concept of an EigenSolver is illustrated in Figure 1. This strategy works with polynomial basis functions such as orthonormal polynomial shape functions formulated for finite element method resulting in the creation of the stiffness matrix first followed by finding the eigenvalue and eigenvector of the thermomechanical propagator matrix extracted from the stiffness matrix. The eigenmodes transform the problem in its frequency domain.

The eigenmodes transform the problem in its frequency domain. Some other methods of interest in this case and which may require further investigation are the Spectral Finite Element Methods (SFEM). Although the method and its variants has not been attempted in this description- the research in this direction has been reviewed.(a)(b)(c)(d)Figure 1: Schematic Diagram illustrating the Concept of an EigenSolver. In (a), first a Gaussian heat flux is applied to the material followed by the split of the flux in orthogonal modes in (b). These orthogonal modes are a function of geometry and the material (in-homogeneity and non- linearity) which traverse forward through the cross-sections shown in (c) followed by their temperature counterparts which traverse backwards towards the source as the flux modes hit the fixed temperature boundary.

In (d), the stiffness modal coupling of flux and temperature modes between changing cross-sections is shown, where black regions represent numbers in a matrix and white regions represent insignificantly small numbers. These matrices will assume an identity matrix or integral multiples of the identity matrix along the matrix diagonal since consecutive cross-sectional material properties/geometry are assembled in a fashion such that they are integral multiples w.r.t one another whereas if the material properties/geometry is continuously changing, this will lead to off-diagonal terms comprising an interaction of one cross-section w.r.t the other.For general wave propagation related problems, the traditional finite element method (FEM) is the most used and trustworthy numerical technique for its capabilities of analyzing complex structures and capturing dynamic response with high accuracy.

However, for reliable predictions of dynamic response of wave propagation, the number of elements required is typically large, which induces enormous computational costs. SFEM uses exponential functions instead of polynomial functions as basis functions and the finite element method is implemented based on variational integration in a frequency-dependent domain. SFEM has been studied in stationary vibrations of fluid-filled pipes, [2] railway car structures [3] and wave propagation in composite tubes[4], composite plates [5, 6] and composite beams [7-9], and diagnosis of crack/damage of structures [10-13]. SFEM has good reliability and efficiency compared to traditional FEM for these types of problems.The Fast Fourier transform (FFT) based spectral finite element (FSFE) method is known for its ability to simulate inverse wave motion problems.

However, due to the Fourier series approximation in the spatial domain and discrete assumption in the temporal domain, the lateral boundary responds with an inaccurate wrap around response for short length waves [14]. The wavelet based spectral finite element (WSFE) method, which is similar to the FSFE method, alleviates this problem by using compactly orthonormal and supported Daubechies scalingfunctions [15] approximation both in time and space. The WSFE method allows imposition of initial values with no periodicity assumption and thus is free from wrap around problems [16]. The wave equations are reduced to the ordinary differential equations (ODEs) in both temporal and spatial domains and are typically decoupled using eigenvalue analysis and solved for obtaining the element shape functions.MethodologyA novel solution strategy for prediction of thermo-mechanical variables away from the point of laser exposure has been formulated.

The involved strategy has been borrowed from structural vibrations and image analysis algorithms where modal contributions (eigenmodes) are computed in terms of orthogonal functions. The typical approach for determination of modes using sine or cosine functions has been extended to fit a finite element framework where an eigenvalue problem of the thermomechanical propagator matrices (these matrices compute the propagation of the thermomechanical field from one cross-sectional layer to another along the –z direction) is for computed thermomechanical fields. It has been observed that very near to the point of laser exposure the number of modes required to reconstruct the solutions is very high (the number of modes is limited by the number of nodal points in each cross-section) and as the distance from the point of laser exposure increases, the number of modes required to accurately predict the solution drops very quickly.

For performing these calculations, therefore, the region very close to the point of laser exposure in the –z direction where non-linear thermal conductivity relations and melt pools are active is solved using a unique Feed Forward Dynamic Adaptive Mesh Refinement and De- Refinement (FFDAMRD) finite element method. When more than 4-9 layers from the top surface, the problem is solved using the thermal eigensolver approach (as illustrated in Figure 2). For performing the calculations on the current time step, first the thermal evolutions in the first 4-9 layers are calculated using FFDAMRD and the thermal distribution of the deeper layers as a boundary condition. This is followed by using the new thermal distribution calculated for the 4th to 9th layers as flux boundary conditions to solve for the deeper layers in the current time step.in orderto determinethe orthogonal basisfunctions(eigenvectors)Figure 2: XZ plane cross-section elements viewed from the -Y direction.

The Solid-Powder pattern in the +X direction is repeated and extruded in the –Z direction, illustrating a part with the same cross-sectional geometry between layers. S denotes the portion of the powder bed which has converted from powder to solid following laser exposure and P denotes the portion of the powder bed which didn’t change its state and remained as powder.In an another development, a new solution methodology has been formulated such that the thermal solution using our FFDAMRD algorithm can be updated in a manner such that LU factorization need not be calculated more than once to obtain the thermal distribution for every time step (shown by red dots) in Figure 3 for each offset location (Figure 3 compared to Figure 4 show how the fine mesh moves by one coarse mesh division in the direction of laser motion, thus defining one offset).

The method used for this development is a multi-rank/order update for changing thermal stiffness as a function of time step/offset.Figure 3: Figure showing the combined fine and coarse mesh at offset 1 with red dots showing the time stepping.Figure 4: Fine mesh movement from offset1 to offset2.Results and DiscussionThermal EigensolverIn the case of Ti6Al4V, accurate thermomechanical solutions can be calculated almost instantaneously 4 or more layers away from the point of laser exposure using a limited number of modes. A part with 2.5 dimensional properties (where the geometry is extruded in the z direction) in the powder bed, due to the scalar nature of eigenmode decay equations, turns out to be extremely uncomplicated, as shown in Figures 5, 6 and 7. The amount of error introduced using this approach is negligibly small for a point energy source (see Figure 5) while the speed improvement is substantial.

The amount of error introduced using this approach is negligibly small for a point energy source (see Figure 5) while the speed improvement is substantial. For cases of a line energy source (see Figure 6), the match is good but not excellent.For cases of a large area source, due to unrefined orthogonal modes, the trend of solutions is replicated but the error is much worse when compared to a point energy-source solution (see Figure 7). The use of Arnoldi and other methods for refinement of modes [17] for this scenario would be relevant for manufacturing processes in which point sources of energy are not used.

The use of Arnoldi and other methods for refinement of modes [17] for this scenario would be relevant for manufacturing processes in which point sources of energy are not used. For focused energy sources such as for SLM and EBM, this approach completely eliminates the need to re- compute thermomechanical fields using fine-scale FFDAMRD for the problem domains which are more than a few layers away from the laser; thus eliminating the vast majority of the computational time necessary to simulate real parts (which are made up of thousands or tens of thousands of layers).Figure 5: Nodal point by point match of a finite element solution for a thermal field with a Modal reconstructed (eigensolver) solution for a Gaussian point energy source. The error is negligibly small (<0.1%) and the results lie right on top of each other.Figure 6: Nodal Point by point comparison of a finite element solution with a Modal reconstructed solution for a constant line energy source.

The match is good but not excellent when compared to Figure 5.Modal Reconstruction100.84000Finite Element Solution100000.250000.4# of nodal pointsLinear thermal field solution (normalized)Area Solution200030000.6Figure 7: Nodal Point by point comparison of a finite element solution with a Modal reconstructed solution for an area energy source. The modally reconstructed solution matches the trend but not the magnitude of the solution due to unrefined orthogonal modes. This result is irrelevant for moving point energy problems associated with SLM or EBM, but should be taken into account when seeking the correct solution for area energy sources.For problems which do not have an extruded geometry from the top surface to the bottom surface the solution becomes intractable since the extruded geometry eigensolutions could nolonger be used for these cases and they require further mathematical corrections to account for the inhomogeneity.

The finite element solution has been superposed on an extruded homogenous eigensolution in Figure 8 for a case shown in Figure 9. It can be clearly seen that there exists a significant mismatch between the finite element and thermal eigensolver results for this case if no correction is implemented.Figure 8: FEA and Thermal eigensolver with 0th order correction for node-by-node comparison using homogenous modes for computationXZ planeSolidSpreadSpreadPenetrationPowderFigure 9: Inhomogeneous case in the -z direction showing a solid part with an internal cavity and entrapped powderUsing a first order correction, the results match with a relative error of 0.14%. FEM results superimposed on the thermal eigensolver results with first order correction are shown in Figure 10.

FEM results superimposed on the thermal eigensolver results with first order correction are shown in Figure 10. The eignesolver solution is ~8000 times faster for this small problem, but.the computational advantage will be even greater for larger problems as the matrix size grows exponentially with number of layers.Figure 10: FEA and Thermal eigensolver with 1st order correction one-on-one DOF comparison using homogenous modes for computation.It should be noted that, so far, eigenvalues have been identified and reconstructed for the thermal case only. A parallel architecture is required for solving the deformation problem using our dislocation density based crystal plasticity (DDCP) finite element solver.

A parallel architecture is required for solving the deformation problem using our dislocation density based crystal plasticity (DDCP) finite element solver. For DDCP, the eigenvalue problem has orders of magnitude higher complexity than for the FFDAMRD thermal problem and will thus require a very fine refinement to compute the dispersion matrix and penetration matrices correctly in order to obtain the propagator matrix for its eigenvalue determination.Multi-rank updatesIn order to use the multi-rank updates for computing the approximate inverse of the changing thermal stiffness as a function of offsets, the convergent singular values have been plotted against the total number of singular values in Figure 11. It is clear from this illustration that the singular values converge as a function of increasing number of mesh transfers.

It is clear from this illustration that the singular values converge as a function of increasing number of mesh transfers. Another important aspect from Figure 11 is that the singular values follow a steeply decreasing curve similar to a rectangular hyperbola and hence the number of singular values required for computing the inverse should be less than the full size of the matrix. This results in even more savings in computational effort and time along with convergence of the singular values as a function of number of offsets.

This results in even more savings in computational effort and time along with convergence of the singular values as a function of number of offsets. A one-to-one match between the thermal distributions obtained using the LUdecomposition and a much lower rank 1000 update of the thermal stiffness (including dynamic C) is shown in Figure 12.Figure 11: Singular values as a function of fine mesh transfer offsetsFigure 12: Thermal distributions computed by performing the LU/Cholesky Decomposition of the full thermal stiffness matrix against a much lower rank update of the inverse.Multi-order updatesFor using the multi-order updates, the inverse of the thermal stiffness for the current time step has been updated using the thermal stiffness of the previous time step. While multi-rank updates were found to be more beneficial and faster in the case of fine mesh offset transfers, multi- order updates are more suitable for updating the thermal distribution as a function of time steps.

A one-to-one thermal DOF comparison using Cholesky factorization in Finite Element and using the 4th order thermal stiffness update is shown in Figure 13. The results are in excellent agreement with the thermal distribution obtained using the finite element method.Figure 13: Thermal distributions computed by performing the LU/Cholesky Decomposition of the full thermal stiffness matrix against a much lower order update of the inverse.Conclusions & Future WorkThree algorithms have been developed to increase the efficiency of computation of thermal fields in metal melting based additive manufacturing. The first algorithm deals with propagation of the macroscopic thermal response in the –z directions (a multi-order update followed by update of the macroscopic 3 dimensional thermal response).

The first algorithm deals with propagation of the macroscopic thermal response in the –z directions (a multi-order update followed by update of the macroscopic 3 dimensional thermal response). The other two algorithms provide two different methods for computing thermal distribution as a function of fine mesh transfer offsets and time steps inside a particular fine mesh offset configuration. The results of these algorithms have been found to be in an excellent agreement with their inefficient finite element counterparts. These algorithms are being implemented in the commercial code being developed by 3DSIM, LLC in close collaboration with University of Louisville personnel.References1.2.3.4.5.6.7.8.9.10.11.Berdichevskii, V., Variational-Asymptotic Method of Constructing a Theory of Shells: PMM vol. 43, no. 4, 1979, pp. 664–687. Journal of Applied Mathematics and Mechanics, 1979. 43(4): p. 711-736.

However, its suitability for powder bed fusion (PBF) is limited due to its susceptibility to thermo-oxidative aging, resulting in material degradation and mechanical property deterioration over time. To address this issue, the efficacy of antioxidants in increasing the aging resistance of PA6 in PBF was investigated. Process-adapted analysis was employed using a coupled rheometer FTIR instrument to elucidate physical and chemical changes in the material. In addition, the viscosity number of the virgin and processed powder, the yellow index, and the part performance were evaluated. Results revealed that the addition of primary and secondary antioxidants significantly enhanced the aging resistance of PA6 in PBF, thereby increasing its potential as a suitable material for additive manufacturing applications.IntroductionPowder bed fusion of polymers (PBF-LB/P) is an additive manufacturing process that uses a laser to selectively melt a powder material.

It creates prototypes and functional parts with complex geometries and high accuracy. A thin layer of powder material is spread out, and a laser traces the desired part shape on the powder bed, melting the powder in that area. The platform is lowered, and another layer of powder is added. This layer- by-layer process is repeated until the parts are built. [1] Various materials are commonly utilized in PBF-LB/P, with polyamide 12 (PA 12) being the most extensively employed material to date. Despite polyamides holding a significant market share in traditional manufacturing processes, PA 12 has a relatively minor role and is primarily used in special applications. The more commonly used polyamides, such as Polyamide 6 (PA 6) and Polyamide 6.6 (PA 6.6), are rarely utilized in PBF-LB/P. Although commercial PA 6 powders are available, challenges persist regarding the mechanical performance and aging behavior of the material.

Although commercial PA 6 powders are available, challenges persist regarding the mechanical performance and aging behavior of the material. [2] Typically, process temperatures in PBF-LB/P fall within the range of 170-300°C. These temperatures are determined by the melting and crystallization points of the polymer. The processing window is defined as the temperature gap between these points. Although the process chamber is filled with an inert gas atmosphere, a residual oxygen content remains. The level of this residual oxygen varies depending on the specific machine. As a consequence of the high temperature, extended process duration, and remaining oxygen, aging effects manifest in the powder material. This leads to reduced reusability and potential alterations in part properties. [1] Stabilizers can be employed to mitigate the aging effects of polymers during processing and application.

This leads to reduced reusability and potential alterations in part properties. [1] Stabilizers can be employed to mitigate the aging effects of polymers during processing and application. Typically, these stabilizers are customized for traditional processes characterized by short processing times and high temperatures, or for parts subjected to low temperatures and extended aging periods [3]. Unfortunately, for boundary conditions in PBF-LB/P, there is little to no knowledge of the behavior of traditional polymer stabilization. Therefore this work will focus on the aging behavior of PA 6 powder, and the stabilization of the material regarding thermo-oxidation effects. Aging effects in polymers can generally be classified into two categories: physical aging and chemical aging. Physical aging occurs when the polymer molecules undergo rearrangements and changes in their physical structure due to external factors such as temperature, humidity, and stress.

Physical aging occurs when the polymer molecules undergo rearrangements and changes in their physical structure due to external factors such as temperature, humidity, and stress. These effects are reversible after the melting of the polymer. Chemical aging occurs when the polymer molecules react with external agents, such as oxygen, UV light, and other chemicals. These reactions can lead to changes in the chemical structure of the58polymer, resulting in changes in its properties, such as color, solubility, and strength. Chemical aging is non- reversible due to changes in the polymer chain. [4] Several chemical aging mechanisms can contribute to changes in polymers. These include thermal degradation, post-condensation, UV degradation, hydrolysis, oxidation, or a combination of these factors. Thermo-oxidation, for instance, can pose significant challenges in specific polymer applications.

Thermo-oxidation, for instance, can pose significant challenges in specific polymer applications. This is particularly true for applications involving high temperatures or prolonged exposure to oxygen. Autooxidation is another process that occurs in polymers upon exposure to oxygen. It results in the formation of free radicals, which can cause chain scission and other forms of degradation. The autooxidation cycle is a series of reactions that can occur in polymers during this process. This cycle is divided into three main stages: [3] Initiation: This represents the initial stage of the cycle, where oxygen molecules interact with the polymer and generate free radicals. Various factors such as heat, UV radiation, and the presence of specific catalysts can trigger this process. Propagation: During this stage, the free radicals formed in the initiation phase react with oxygen to produce new free radicals.

Propagation: During this stage, the free radicals formed in the initiation phase react with oxygen to produce new free radicals. These new radicals then further react with other polymer molecules, initiating a chain reaction that can result in polymer degradation, branching, or chain-extantion. Termination: The final stage of the cycle involves the reaction of free radicals with one another, forming stable products. This termination step brings an end to the chain reaction.In PBF-LB/P studies mainly focus on the chemical aging by means of post-condensation of end groups unregulated polyamides such as PA12 [5] or PA11. During the process, chemical aging occurs in the PA 12 powder. In other studies, this material change is attributed to post-condensation in the solid phase [6]. The chain ends of the PA 12 macromolecules can react with each other, releasing water [7].

In other studies, this material change is attributed to post-condensation in the solid phase [6]. The chain ends of the PA 12 macromolecules can react with each other, releasing water [7]. This results in an increase in molecular weight [6,8], and a corresponding increase in the melt viscosity of the material [9]. Changes in the average molecular weight are characterized, for example, using the viscosity number (VZ) or the melt volume flow rate (MVR) [6,10]. In addition to chemical aging, physical aging of the macroscopic particle morphology occurs at the particle level. Due to the tribo-mechanical stress during powder application and the accompanying thermal stress during the recoating process, the flowability of the powder decreases [6]. This decrease in powder flowability has been attributed to the smoothing of the particle surface, and a change in the concentration of flow aids on the particle surface [6].

This decrease in powder flowability has been attributed to the smoothing of the particle surface, and a change in the concentration of flow aids on the particle surface [6]. This results in a decrease in the distance between the individual particles and an increase in the adhesion forces. However, no change in the particle size distribution was observed [6,10]. Futher, there is a lot of work on thermo-oxidative aging and stabilizers, often with the focus on component properties after certain time intervals. Octavie et al. [11] investigated the aging for PA 11 for injection moulding. Regarding the impact of temperature during tensile testing, it has been noted that higher temperatures can delay the process of embrittlement. In other words, the decrease in strain at break occurs at higher oxidation conversion degrees.

In other words, the decrease in strain at break occurs at higher oxidation conversion degrees. The origin of the transition from ductile to brittle is well known within a certain range of molecular weights [12], and allows the definition of a critical molar mass (M'c). When the average molar mass (Mn) exceeds the M'c, the sample exhibits ductile behavior. Other investigations have been carried out regarding the thermal oxidative aging of different polyamide variants. One such example involves the use of FT-IR spectroscopy by researchers to analyze the chemical and physical transformations that take place in PA 6 and PA 6.6 due to thermal degradation. The IR spectra obtained from these studies demonstrates comparable modifications in both PA 6 and PA 6.6, including the creation of carbonyl groups as a consequence of oxidation processes [13]. To facilitate the design of materials for the PBF-LB/P process, a comprehensive understanding of polymer aging mechanisms is crucial.

To facilitate the design of materials for the PBF-LB/P process, a comprehensive understanding of polymer aging mechanisms is crucial. While it is not possible to stabilize a polymer against thermal degradation, here are several strategies to mitigate the effects of oxidation. These include the use of antioxidants, stabilizers, and minimizing the polymer's exposure to oxygen and other degradation sources [3]. Stabilizers are substances incorporated into polymers to protect them from degradation caused by factors like heat, UV radiation, and oxygen. Various types of stabilizers are employed to safeguard polyamides from degradation, including antioxidants, UV absorbers, heat stabilizers, and processing stabilizers. Antioxidants are classified into primary and secondary antioxidants [3].59Primary antioxidants function by reacting with oxygen-containing polymer radicals formed during autooxidation.

Antioxidants are classified into primary and secondary antioxidants [3].59Primary antioxidants function by reacting with oxygen-containing polymer radicals formed during autooxidation. These highly reactive radicals are converted into inactive or less reactive compounds, effectively interrupting the radical chain reaction. Consequently, the antioxidants are consumed in this process [3]. Secondary antioxidants, on the other hand, are capable of reacting with hydroperoxides and decomposing them into non-radical products. By working in tandem with primary antioxidants, secondary antioxidants prolong the life span of primary antioxidants by decomposing hydroperoxides into non-radical compounds, thus impeding the formation of new radicals.

The complementary actions of different types of stabilizers provide enhanced protection against excessive consumption [3].Materials and MethodologyMaterial and machineThe experimental investigations employed a PA 6 material known as Ultramid B27 E, manufactured by BASF in Germany. This specific grade of Ultramid B27 E possesses a low viscosity, rendering it highly suitable for compounding and the production of monofilaments. Three stabilizers were used to increase the aging resistance: Two primary and one secondary antioxidant. Sterically hindered phenolic antioxidantPhenolic antioxidants are radical scavengers that prevent thermal degradation of many organic materials, such as polymers [14]. Their effectiveness increases when used in combination with phosphites, thioethers, and light stabilizers. Polymer radicals that are generated are effectively quenched through hydrogenation by the hydroxyl group. Therefore, phenolic antioxidants act as primary antioxidants.

Polymer radicals that are generated are effectively quenched through hydrogenation by the hydroxyl group. Therefore, phenolic antioxidants act as primary antioxidants. OrganophosphateOrganophosphate antioxidants are widely used additives in polymers to protect against oxidative degradation. They function by intercepting and neutralizing free radicals that initiate and propagate the oxidation process [15]. They act as secondary antioxidants, and efficiently convert hydroperoxides into non-radical products through their function as oxygen-scavenging phosphites. The usage of organophosphate antioxidants in combination with phenolic antioxidants can result in synergistic effects.The materials were dried according to the manufacturer's recommendation and compounded into the PA 6 matrix using a twin-screw extruder MIC 27GL/44D from Leistritz GmbH, according to Table 1. Finally, the manufactured granulate was milled cryogenically to a powder material.

Finally, the manufactured granulate was milled cryogenically to a powder material. The powder was dry blended with 0.1 wt.-% hydrophobic silica and 0.1 wt.-% pyrogenic alumina to improve the powder flowability.Table 1: Compound recipes and additive concentration in wt.-%organophosphate sterically hindered phenolic antioxidant – Type A sterically hindered phenolic antioxidant – Type BReference (Ref.)0 % 0 % 0 %Compound 1 (D1) 0.3 % 0.1 % 0 %Compound 2 (D2) 0.3 % 0.0 % 0.1 %The experimental procedures were conducted utilizing a custom build laser sintering system capable of flexibly customizing all processing parameters. This system incorporates eight quartz radiator arrays, each independently controlled to ensure a homogeneous temperature distribution across the surface of the powder bed. A CO2 laser, possessing a maximum nominal power of 50 W and a spot diameter of 500 µm, was employed to selectively melt the powder material.

A CO2 laser, possessing a maximum nominal power of 50 W and a spot diameter of 500 µm, was employed to selectively melt the powder material. To minimize powder consumption, a downsized build chamber with dimensions of 100 x 100 mm² was utilized. Following a duration of one hour, the machine successfully achieved a residual oxygen content of approximately 1.5% measured inside the process chamber with an oxygen meter. Prior to the processing stage, the powder underwent a drying process for a duration of four hours at a temperature of 70 °C in a vacuum atmosphere. The build procedure consisted of two layers, from which tensile bars were manufactured following the specifications outlined in [16]. A total of four tensile bars in xy-orientation were produced in each layer, employing an energy density of 0.06 J/mm². To achieve the desired energy density, a laser power of 18.7 W, a scan speed of 1250 mm/s, and a hatch distance of 0.25 mm were employed.

To achieve the desired energy density, a laser power of 18.7 W, a scan speed of 1250 mm/s, and a hatch distance of 0.25 mm were employed. The build chamber temperature was set at 210 °C, while the removal chamber temperature was maintained at 150 °C. Each layer had a thickness of 120 µm, resulting in a layer time of 30 seconds. Prior to the start of the build60process, the machine was pre-heated for two hours, initially having a powder height of 5 mm. During this pre- heating period, an additional 5 mm of powder was added. The total build time amounted to 1.5 hours, after which the job was cooled for a duration of four hours within the machine.Analysis MethodsParticle size distributionThe powder particle size distribution (PSD) was determined by image analysis using a QICPIC (SympatecGMBH). In total 1,000,000 particles were used to calculate the PSD. Compression depthThe powder compressibility was measured using the DHR 2 rotational rheometer from TA Instruments [11].

In total 1,000,000 particles were used to calculate the PSD. Compression depthThe powder compressibility was measured using the DHR 2 rotational rheometer from TA Instruments [11]. The measurements were conducted using a setup consisting of 25 mm parallel plates. The bottom plate was modified with a ring of an inner diameter of 27 mm to contain the powder material. To ensure consistency, a normal force of 0.2 N was applied to the sample before it was compressed at a constant rate by decreasing the gap height of the rheometer. The normal force was analyzed during the process. The measurements were repeated ten times. DSC-MeasurementsDynamic and isothermal measurements were performed with the virgin powder by differential scanning calorimetry (Discovery 2500, TA-Instruments) under a dry N2 atmosphere to characterize the melting and crystallization behavior of the powders.

For the dynamic measurements, (2.5 ± 0.1) mg samples were heated and cooled twice from 0 °C to 260 °C with a 10 K/min ramp. For isothermal measurements, samples weighing (2.5 ± 0.1) mg were heated to 250 °C with a 60 K/min ramp and held for one minute. Subsequently, the samples were cooled down to a temperature of 200 °C at a rate of 60 K/min and kept isothermal for two hours. Rheology + FTIRIn this study a combined measurement instrument called the rheonaut, manufactured by Thermo Electron GmbH in Germany, to measure the time depending aging effects simultaneously is used. The rheonaut integrates two techniques: Fourier-transform infrared spectroscopy (FTIR) and a rheometer as presented in [17,18]. FTIR was employed to evaluate changes in the polymer chain. This technique involves directing an infrared beam through an attenuated total reflectance (ATR) crystal into the sample, which is then analyzed using a deuterated triglycine sulfate (DTGS) detector.

This technique involves directing an infrared beam through an attenuated total reflectance (ATR) crystal into the sample, which is then analyzed using a deuterated triglycine sulfate (DTGS) detector. The FTIR measurements were performed using a Thermo Fisher Scientific Nicolet IS10 instrument; absorption measurements were recorded with a spectral resolution of 4 cm⁻¹ and 16 repeat scans. A spectral image was recorded every minute throughout the measurement. Simultaneously, the material's rheological properties were measured using a rheometer, specifically the HAAKE MARS 60 model. The rheological measurements were carried out with the following parameters: a frequency sweep with an angular frequency ranging from 0.1 rad/s to 628 rad/s was performed every 5 minutes. The gap between the parallel plates was set at 1 mm, and the temperature was set at 230 °C. A strain of 1% was applied during the measurements, and the experiments were conducted in ambient air.

The gap between the parallel plates was set at 1 mm, and the temperature was set at 230 °C. A strain of 1% was applied during the measurements, and the experiments were conducted in ambient air. The rheometer utilized a 25 mm parallel plate measurement setup, with the ATR crystal positioned on the boundary area of the 25 mm plate, still covered with polymer. The measurement was carried out for three hours for each sample. Viscosity numberThe viscosity number (VN) is determined to characterize the changes in the viscosity and thus average molecular weight due to aging. The VN is measured according to DIN EN ISO 307. Sulfuric acid at 25 °C was used as the solvent due to security and health issues with a polymer concentration of 0,005 g/cm³. Part and mechanical propertiesThe part thickness and width were measured by a micrometer at three positions. The density was calculated by the part volume and weight. A precision scale was used to determine the part weight.

The density was calculated by the part volume and weight. A precision scale was used to determine the part weight. The surface roughness was measured using a tactile stylus instrument Waveline 20, with a stylus tip of 2 μm and an attaching force of 0.8 mN. The roughness Ra was determined according to [19]. The specimens were conditioned in a vacuum after removing from the part cake leading to a dry state. The mechanical properties were measured following [20] with a speed of 0.5 mm/min for the Young's modulus and 2.5 mm/min for the elongation at break with a testing machine of the type 1484 from Zwick Roell.61Results and DiscussionFor a detailed understanding of the aging effects occurring and the stability of the manufactured powders, time-dependent data is essential. Within the PBF-LB/P process, it is impossible to generate powder samples with a defined thermal history, atmospheric conditions, and aging time.

Within the PBF-LB/P process, it is impossible to generate powder samples with a defined thermal history, atmospheric conditions, and aging time. Therefore, experiments with well-defined boundary conditions are conducted. Structural changes in the polymer chain can influence the melt rheology, with chain scisson leading to a decrease in viscosity and chain elongation increasing viscosity. The frequency- dependent rheological behavior is to demonstrate structural changes in the polymer. The viscosity of the reference material (Figure 1, left) shows no change up to 109 minutes. After this time, the viscosity increases at low frequencies, progressively increasing with time. This increase can indicate cross-linking. D1 (Figure 1, middle) exhibits good aging stability with little to no change over the measured time. D2 (Figure 1, right) exhibits a similar increase in viscosity at low frequencies, starting at 109 minutes.

D2 (Figure 1, right) exhibits a similar increase in viscosity at low frequencies, starting at 109 minutes. Overall, both D1 and D2 illustrate an increased viscosity compared to the reference, suggesting that the antioxidants are working already in the powder manufacturing process.

Furthermore, the reference material shows an increase in viscosity for all measurements at frequencies lower than 2 rad/s, indicating aging effects occurring in the powder manufacturing process.s * a P niy t i s o c s vi105104103time in min:14 26 38 50 62 73 85 97109 121 133 145 156 168 180 192s * a P niy t i s o c s vi105104103time in min:14 26 38 50 62 73 85 97109 121 133 145 156 168 180 192s * a P niy t i s o c s vi105104103time in min:14 26 38 50 62 73 85 97109 121 133 145 156 168 180 192x e p m o cl102x e p m o cl102x e p m o cl10210110110110-1102 101 100 frequency in rad/s10310-1102 101 100 frequency in rad/s10310-1102 101 100 frequency in rad/s103Figure 1: Rheological aging behavior under defined boundary conditions, Ref. (left), D1 (middle), D2 (right), T = 230 °C, air atmosphereThe FTIR analyses provide a deeper understanding of occurring aging mechanisms.

(left), D1 (middle), D2 (right), T = 230 °C, air atmosphereThe FTIR analyses provide a deeper understanding of occurring aging mechanisms. Figure 2 (left) displays the full spectrum of the FTIR for the reference material, with each line representing a spectrum taken at a 1- minute delay. Important peaks are marked to enhance understanding. For oxidation, the range between 1700 and 1800 cm-1 (Figure 2, middle) is of particular significance, where changes associated with degradation products resulting from oxidation are observed. Examples of such changes include the presence of carboxyl or ester groups. In the reference material, the formation of these peaks is observed over time. Figure 2 (right) shows the peak changes (1716 cm-1, kentone) for the reference material, D1, and D2. While the reference material shows an immediate increase, D1, and D2 exhibit stable peaks for up to 30 minutes. After 30 minutes, the peak of D2 starts to increase, while D1 remains stable up to 60 minutes.

The measurements conducted in this study indicate that the additives used did not have a significant influence on these factors (Figure 3, left). The d50.3 is approximately 80 µm, with a slightly increased fraction of smaller particles. The largest particles observed are up to 200 µm. The PSD is considered suitable for the process. SEM analysis reveales the presence of typical edgy particles, which are commonly observed in cryogenic milling processes (Figure 3, middle). Edgy particles can lead to bad recoating behavior. The powders without any additives are shown to exhibit a rough and cracked powder bed (Figure 3, right). However, the use of flow agents can enhance powder flowability, resulting in a smoother powder bed. The addition of hydrophobic silica increases the distance between particles and reduces adhesion forces, promoting better particle flow. Pyrogenic alumina acts as an antistatic agent, preventing the buildup of static charges.

Figure 4 (left) shows the dynamic DSC-measurements of the three powder materials. For the melting peak of the first heating, no significant difference is detected. For the first cooling, a slight shift of the crystallization peak to lower temperatures is observed for D1 and D2 compared to the reference powder. This indicates that the used antioxidants have no nucleating effect. The measured processing windows for all three materials are shown in Figure 4 (middle) with a slightly wider processing window for D1 and D2. While the dynamic DSC is an indicator of the processing temperature and the processing window, it is hard to transfer the results to the PBF-LB/P process. In process, the thermal boundary conditions differ strongly from the 10 K/min heating and cooling rate. Therefore, Figure 4 (right) displays isothermal measurements for the virgin powders.

In process, the thermal boundary conditions differ strongly from the 10 K/min heating and cooling rate. Therefore, Figure 4 (right) displays isothermal measurements for the virgin powders. Compared to the reference, the crystallization peak of the D1 and D2 powder is shifted to shorter times, indicating a nucleation effect. These results indicate that process adapted measurements like isothermal DSC-measurements are important to understand the material behavior in the process. A shorter isothermal crystallization time can lead to warpage or process termination due to curling.631.0Ref. D1 D22.5D22.00.8g / W ng / W n1.51.50 min0.61.0iiw o l f1.79 minD1w o l f0.50.43.63 mint a e ht a e h0.00.50.2Ref.

D1 D22.5D22.00.8g / W ng / W n1.51.50 min0.61.0iiw o l f1.79 minD1w o l f0.50.43.63 mint a e ht a e h0.00.50.2Ref. D1 D2Ref.1.0processing window0.01.5200 190 temperature in °C Figure 4: Dynamic DSC-Measurements of the three materials (left), measured processing window (middle), and isothermal Measurements of the virgin1802102200510 time in min1520100 125 150 175 200 225 250 temperature in °Cexo up1.00.80.60.40.20.0powder, Tiso = 200 °C (right)Due to thermo-mechanical stress in the recoating process, the powder flowability is reduced over the process. This physical aging mechanism can lead to smoothing of the particle surface or flow additives which are pressed into the material. Both lead to a decrease of the particle – particle distance and leads to a decreased powder flowability.

Both lead to a decrease of the particle – particle distance and leads to a decreased powder flowability. Figure 5 (left) shows the measured compression depths for the three powders in virgin state and after one processing cycle. In the virgin state, the powders indicate a similar compression depth. After processing, a slight increase for the D1 and D2 powder is measured compared to the reference material. This may lead to problems in recoating or a decrease of the powder bed density. Overall, D1 and D2 show only a small increase which should not affect the process significantly. To analyze the chemical aging, the viscosity number (VN) of the virgin and used powder is measured (Figure 5, middle). The reference powder in the virgin state shows an already decreased VN compared to D1 and D2. The same behavior was analyzed in the rheological measurements of the viscosity (Figure 1Figure 5). This suggests a stabilising effect of the used antioxidants in the powder manufacturing process.

The same behavior was analyzed in the rheological measurements of the viscosity (Figure 1Figure 5). This suggests a stabilising effect of the used antioxidants in the powder manufacturing process. After processing, the VN decreases strongly for all materials. This indicates chain scissoring due to thermal oxidation. Due to the reactive end groups, polyamides tend to increase in molecular weight (post-condensation). Figure 5 (right) details the change of the VN.

Although this change does not directly correlate with the part properties, a discolored part cannot be sold. Figure 6 (left) displays the virgin powder, used powder, and the produced part. Slight discoloration is observed for D1 and the reference material compared to D2 in the virgin powder; this indicates the stabilizing effect during the powder manufacturing process. After one cycle, both the powder and parts exhibit significant discoloration for all three materials. The reference material shows the strongest color64n in o i t a zi l la t s y r cf o e e r g e dchange, severely limiting the reuse of the powder, especially for larger build jobs where this effect is expected to amplify. Figure 6 (right) presents the measured yellowness index, where the reference material exhibits the most pronounced changes.

In comparison, D1 (Figure 7, middle) and D2 (Figure 7, right) show no pores, good layer bonding, minimal presence of non-molten particles, and well- defined geometrical shapes. The inferior geometry of the reference material can be attributed to its decreased viscosity and VN.Figure 7: Thin cuts obtrained from the center of the manufactured tensile bars, reference material (left), D1 (middle), D2 (right)The presence of the shown pores leads to a decrease in part density for the reference material (Figure 8, left). The density increases for D1 and reaches its highest value for D2. While the thin cuts only indicate the local position of the tensile bar, the density considers the entire bar. Pores located in the shoulder or other positions of the tensile bar can reduce the part density, suggesting that D2 exhibits the highest density and the lowest pore volume. In addition to density, the surface roughness of the part plays a critical role in its mechanical properties.

In addition to density, the surface roughness of the part plays a critical role in its mechanical properties. D2 exhibits a slightly increased surface roughness compared to the reference material and D1. The surface roughness is primarily influenced by the particle system used and the presence of partially molten particles on the surface area. The measured E-Modulus shows an increase for the D1 and D2 materials compared to the reference material (Figure 8, middle). Material properties can be influenced by aging effects, which in turn affect the E-Modulus. Additionally, the geometry and presence of pores can result in an inaccurate measurement of the cross-section, further contributing to the lower E-Modulus observed in the reference material. Regarding elongation at break, a significant increase is observed for D1 and D2 compared to the reference material (Figure 8, right).

The experiments conducted in the study include particle size distribution, compression depth, DSC measurements, rheology and FTIR, viscosity number, and part and mechanical properties. The study used a combined measurement instrument called the rheonaut to measure the time-dependent aging effects under defined boundary conditions. The results revealed that the viscosity of the reference material showed no change up to 109 minutes, after which it increased at low frequencies, indicating cross-linking due to aging effects. The FTIR measurements provided a deeper understanding of the aging mechanisms at play. The results showed that the reference material exhibited an immediate increase in the formation of peaks associated with degradation products resulting from oxidation, such as the presence of carboxyl or ester groups. However, D1 and D2 exhibited stable peaks for up to 30 and 60 minutes, respectively, indicating the stabilizing effect of the used antioxdants.

However, D1 and D2 exhibited stable peaks for up to 30 and 60 minutes, respectively, indicating the stabilizing effect of the used antioxdants. The particle size distribution was determined using image analysis, and the results indicated that the additives used did not significantly PSD or the recoating behavior. All grinded powder materials needed additional flow additives for a suitable recoating behavior. The compression depth was measured using a rotational rheometer, and the results showed a slight increase in compression depth for D1 and D2 powders after processing, which could potentially lead to recoating problems or a decrease in powder bed density. Further, dynamic and isothermal measurements were performed to characterize the melting and crystallization behavior of the powders. The results showed that the used antioxidants did have a nucleating effect for the isothermal crystallization. The dynamic measurements showed slightly wider processing window for D1 and D2.

The results showed that the used antioxidants did have a nucleating effect for the isothermal crystallization. The dynamic measurements showed slightly wider processing window for D1 and D2. The viscosity number was determined to characterize the changes in viscosity and average molecular weight due to aging in the process. The results showed that the viscosity number decreased strongly for all materials after processing, indicating chain scissoring due to thermal oxidation. D2 showed the smallest change and, therefore, the best aging stability for the used powders. The study also examined the discoloration of the polymer due to thermo-oxidation, which, while not directly affecting the part properties, can impact the marketability of the product. The results showed that slight discoloration could be observed for D1 and the reference material compared to D2 in the virgin powder, indicating the stabilizing effect during the powder manufacturing process.

After one cycle, both the powder and parts exhibited significant discoloration for all three materials. The reference material showed the strongest color change, severely limiting the reuse of the powder, especially for larger build jobs where this effect is expected to amplify. However, the use of antioxidants demonstrated a reduction in the discoloration of polyamide 6. The yellowness index measurements confirmed these observations, with the reference material exhibiting the most pronounced changes. The manufactured tensile bars show that D2 exhibited the highest density and the lowest pore volume. Further, the mechanical properties were measured, and the results showed an increase in the E-Modulus for D1 and D2 materials compared to the reference material. Additionally, a significant increase in elongation at break was66D2observed for D1 and D2 compared to the reference material.

Additionally, a significant increase in elongation at break was66D2observed for D1 and D2 compared to the reference material. In conclusion, the study found that the use of antioxidants can improve the mechanical properties of the manufactured parts while also reducing thermo- oxidation of the material.Acknowledgment This study is funded within the IGF project (21598 N) with the title "Improving the aging stability of plastic powders for selective laser sintering by tailor-made additive systems". The authors would like to thank Niels Brauch of the Fraunhofer Institute for Operational Stability and System Reliability (LBF) in Darmstadt for material selection and compounding. In addition, the authors would like to thank Florentin Tischer of the Institute of Particle Technology at the FAU Erlangen-Nürnberg for providing particle size distribution data.References [1] M.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEnvironmental Effects on the Stress Corrosion Cracking Behavior of an Additively Manufactured Stainless SteelJonathan Pegues2,3, Michael Roach1, R. Scott Williamson1, Nima Shamsaei2,3,*1Department of Biomedical Materials Science, University of Mississippi Medical Center, Jackson, MS 39216 2Department of Mechanical Engineering, Auburn University, Auburn, AL 368493National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849Corresponding author: shamsaei@auburn.eduAbstractAdditive manufacturing (AM) is becoming a more viable manufacturing process in the biomedical, aerospace, nuclear, and defense sectors as a means to fabricate near net shaped parts on demand.

Austenitic stainless steels are gaining attention for AM of products for these applications due to their ease of fabrication, excellent corrosion resistance, and superior toughness. The performance of these alloys fabricated by AM techniques such as laser beam powder bed fusion has not been yet fully established. This research compares the microstructural characteristics and tensile stress corrosion cracking (SCC) of 316L stainless steel at room temperature and physiological temperature in distilled H2O, salt water, Ringers, and spiked pH Ringers conditions.Keywords: Stress corrosion cracking, environmentally assisted cracking, austenitic stainless steel, tensile propertiesIntroductionAdditive manufacturing is rapidly becoming a viable option for parts used across several sectors including biomedical and defense. In both of these applications, the part’s resistance to environmentally assisted cracking (EAC) in chloride environments could be of critical concern.

In both of these applications, the part’s resistance to environmentally assisted cracking (EAC) in chloride environments could be of critical concern. For biomedical applications, surgical implants can be subjected to corrosive attack particularly in the event of an infection where the bodies pH level can drop as low as five [1]. For defense applications, parts are often subjected to sea spray conditions across a broad range of temperatures depending on the location and time of year [2, 3].It has been reported that stress corrosion cracking (SCC) could occur in any alloy system given non-optimal process conditions and environment [2]. Due to the layer-by-layer manufacturing and unique microstructure provided by AM techniques, the susceptibility of these455materials to SCC is unknown.

Due to the layer-by-layer manufacturing and unique microstructure provided by AM techniques, the susceptibility of these455materials to SCC is unknown. While many studies have investigated the mechanical performance of biomedical/defense grade additive manufactured materials such as 316L stainless steel (SS) [4- 7], systematic investigation into the stress corrosion cracking behavior are lacking in the literature.In this study the susceptibility of laser powder bed fusion (LB-PBF) 316L SS to stress corrosion cracking (SCC) under slow strain rate (SSR) testing according to ASTM G129 [8] is investigated and compared to its wrought counterpart. Tests are performed in distilled H2O, Ringers solution, a pH spiked Ringers solution, and a 3.5% solution of NaCl.

Tests are performed in distilled H2O, Ringers solution, a pH spiked Ringers solution, and a 3.5% solution of NaCl. The ductility, strength, and time to failure are all investigated to assess the effect of each environment on the LB-PBF material for both room temperature (RT) and 37° C.Experimental Set-up316L SS rods were fabricated horizontally on an EOS M290 using OEM recommended parameters. Before removal from the substrate, the parts were stress relieved at 400°C for 1 hr and furnace cooled to RT. Threaded grip test coupons were designed based on ASTM G129 and machined to the final dimensions shown in Figure 1.1·IC: 112•-f 2 = 012 I _j69.87 - \ -.39.87 Dimensions are in mm15Figure 1: Specimen geometry used for slow strain rate testing.Microstructural analysis was performed using both optical and scanning electron microscopy (SEM).

Samples taken from the fabricated rods were polished using standard metallurgical procedures and electrolytically etched using a solution of nitric acid. Additional microstructure analysis was performed using electron backscatter diffraction (EBSD).The SSR tensile tests were performed on an MTS servo-hydraulic load frame with a 100 kN capacity at a displacement stroke rate of 2.5*10-5 mm/s which resulted in an approximate strain rate of 1*10-6 s-1. Tests were conducted either at RT or 37°C. Tests conducted in both Ringers and spiked pH Ringers were done only at 37°C while the distilled H2O and 3.5% NaCl were conducted at both RT and 37°C. Test specimens were completely submerged in the test solution and the temperature was monitored and controlled using a Digi-Sense temperature controller. Figure 2 gives a representation of the test setup used for this study.

Figure 2 gives a representation of the test setup used for this study. All samples were tested until failure,456after which, a measuring microscope was used to determine the elongation to failure and reduction of area (RoA).Fracture surfaces were then examined with both a Keyence digital microscope and a Zeiss scanning electron microscope (SEM). Ratios between the control environment and the harsh environment were calculated by dividing the mean values of samples in harsh environments by the mean values of the samples in the control (distilled H2O) at room temperature. ASTM G129 [8] states that ratios that are less than unity are considered to be susceptible to SCC.

ASTM G129 [8] states that ratios that are less than unity are considered to be susceptible to SCC. For this study a more strict definition of ratios falling below 0.9, which has been used in previous studies [9, 10], is considered to be susceptible to SCC.Figure 2: Test set up showing specimen submerged in solution during testing.ResultsThe microstructure of the LB-PBF 316L SS showed strikingly different morphology compared to the wrought counterpart, as shown in Figure 3. The microstructure of the wrought material is characterized by approximately 13 μ m equiaxed grains with a high density of annealing twins which can be recognized by the specific misorientation across the grain boundary. Σ 3 twin boundaries are low energy, mostly coherent grain boundaries which have a specific misorientation of 60° which can vary a few degrees by the Brandon criteria. The LB-PBF material, on the other hand, shows approximately 7 μ m non-equiaxed grains and a low density of annealing twins.

The LB-PBF material, on the other hand, shows approximately 7 μ m non-equiaxed grains and a low density of annealing twins. In457contrast to the wrought material, the LB-PBF has a much higher density of low angle grain boundaries. Porosity was also observed for the LB-PBF material while unsurprisingly being absent from the wrought material.Additive Manufactured 3 l 6L Stainless Steel0.50. 125 0.4 ·-= ~ fi; 0.3... i 0.2Z 0.1s:: 0.10 0 ·-= ~ 0.08 ri; .., 0.06 u.&> § 0.04 Z 0.02w w ~ ~ ~ roro ~ w ~1020~ 40Grain Sile (Diomete<) (microns)Mlsorlentatlon Angle (degr""5]Wrought 3 l 6L Stainless Steel0.140.4= 0 ·-= " .. ri; ... u l0.30.20.1w w ~ ~ ~ ro Grain Size (Oiametet) (microns)ro1020~~Misorlentatlon Angle (degrees)Figure 3: Microstructure of additive manufactured and wrought 316L stainless steel.The stress-displacement curves for LB-PBF and wrought material in distilled H2O at both RT and 37°C is given in Figure 4.

The LB-PBF condition showed higher yield stress (σ y) and ultimate tensile stress (UTS) accompanied by lower elongation to failure. The improved strength of the LB-PBF material in most likely a result of the finer microstructure, i.e. grain boundary strengthening.700600~500 ~ 400 '-" i2 30011) I-<LB-PBF - H20 23 Cczi 200100LB-PBF - H20 37 C - - Wrought - H20 23 CWrought - H20 37 C002468101214Displacement (mm)Figure 4: Stress-displacement curves for LB-PBF and wrought material at both room temperature and 37°C.458~~roroThe fracture surfaces of the failed specimens gave no indication of brittle fracture for any of the test environments. Figure 5 gives a representation of select environments and temperatures. While all environments and temperature combinations are not covered in this image, the results show little to no indication of brittle failure.

While all environments and temperature combinations are not covered in this image, the results show little to no indication of brittle failure. In both wrought and LB-PBF and all environmental conditions, the typical fibrous and ductile shear failure areas were observed. The LB-PBF material did show larger void growth which can be seen as the porous looking areas on the fracture surface. This is most likely due to the presence of the process induced porosity and could be leading to the lower elongation to failure observed for the LB-PBF material.AM - D Water - 37CAM - Ringer's - 37CWrought - D Water - 23CAM - 3.5% NaCl - 37CAM - Ringer's - pH 2 - 37CWrought - 3.5% NaCl - 23CFigure 5: Fracture surfaces of select specimens in various environments and temperatures.The SSR SCC tensile results for both AM and Wrought in all environmental conditions and temperatures are shown in Figure 6a-d.

Similar to the stress-displacement curves shown in Figure 6a, the elongation to failure is lower for all AM tests regardless of environment with the 23°C tests showing slightly higher elongation to failure than the 37°C tests. Interestingly, while elongation to failure and reduction of area (ROA) are both measures of ductility, the ROA did not show the same temperature dependence that was observed for elongation to failure, as shown in Figure 6b. The difference between the AM and wrought material, however, was still evident. Time to failure (TTF) for each condition is shown in Figure 6c and, following the observations made for elongation to failure, the AM material showed much lower TTF than the wrought material with the 23°C showing slightly higher TTF than the 37°C tests. Lastly, the UTS for each condition is shown in Figure 6d.

.•65Sample - Solution - Temp(c)180! . ,i: ll I170160150140130120110100•. .. • •I. •Sample - Solution - Temp(d)700t:.! ..~650600550. a;I500Sample - Solution - Temp460aTo assess the susceptibility of EAC under SSR testing, ASTM G129 requires the comparison SSR ratios of the chloride environment to the control. Decreasing SSR ratios can be indicative of susceptibility to stress corrosion cracking with ratios that drop below 0.9 showing a higher sensitivity to the harsh environment. Table 1 list all ratios for each condition compared to its control. It is important to note that the condition can only be compared to a control with similar conditions. For example, the AM test conducted at 23°C in 3.5% cannot be compared to the control (distilled H2O) for AM conducted at 37°C or with the control for wrought at 23°C.

For example, the AM test conducted at 23°C in 3.5% cannot be compared to the control (distilled H2O) for AM conducted at 37°C or with the control for wrought at 23°C. Surprisingly, Table 1 shows that the AM condition shows very little susceptibility to chloride environments as all ratios are greater than or equal to 0.95.

Surprisingly, Table 1 shows that the AM condition shows very little susceptibility to chloride environments as all ratios are greater than or equal to 0.95. This is similar to what was shown for the wrought condition at 23°C in 3.5% NaCl.Table 1: Failure ratios for all test conditions.Sample TypeTest Temp.Test SolutionElongation to Failure Ratio (ELR)Reduction of Area Ratio (ROAR)Time to Failure Ratio (TTFR)AM Machined23CD Water1.001.001.0023C3.5% NaCl1.020.981.1237CD Water1.001.001.0037CRinger’s1.000.950.9537CRinger’s – pH 21.031.000.9837C3.5% NaCl1.050.971.04Wrought Machined23CD Water1.001.001.0023C3.5% NaCl0.971.011.03ConclusionsResults from this preliminary study indicate that the LB-PBF material does not show susceptibility to SCC under the tested conditions.

While the tensile behavior of the LB-PBF material shows large differences compared to the wrought material, there was little to no evidence of brittle fracture on the fracture surfaces related to EAC. Additionally, ratios of the tensile properties compared to the control were essentially unity, indicating that SCC behavior was significantly different between the control and harsh environment. These results, while positive,461show the promising potential additive manufactured austenitic stainless steels in harsh environments. It must be noted, however, that while slow strain rate testing can be useful in determining the susceptibility of materials to SCC, the service life of these parts are often vastly longer than what takes place during this test method.

Additionally, the fatigue properties of materials used in harsh environments is often more sensitive and, therefore, more rigorous studies are needed to fully understand the effect of EAC on these materials.AcknowledgementsPartial funding for this work was provided by the National Science Foundation under Grant No. 1657195. Supports from the Naval Air Systems Command (NAVAIR) is also greatly appreciated.References[1] M. Sivakumar, S. Rajeswari, Investigation of failures in stainless steel orthopaedic implant devices: pit- induced stress corrosion cracking, Journal of Materials Science Letters 11(15) (1992) 1039-1042.[2] B.F. Brown, Stress-corrosion cracking in high strength steels and in titanium and aluminum alloys, Naval Research Lab Washington DC, 1972.[3] R.P. Wei, R.P.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEqualization of Metal Powder for Binder Jet Printing Michelle Niu, Krzysztof Nauka HP Inc, 1501 Page Mill Rd., Palo Alto, CA 94304AbstractMetal powders used in additive manufacturing often face the issue of variable powder surface characteristics, causing inconsistent part quality. As such, a ”homogenization” treatment for metal powder surfaces may be advantageous for relevant fabrication processes. By using high-powered xenon flash lamp to rapidly heat powder samples, the surface energy of the powder particles was brought to the same value regardless of the powder source and past usage. Pulse-heating the powder caused the atomic structure of particle surfaces to undergo rearrangement and removed adsorbed moieties.

Pulse-heating the powder caused the atomic structure of particle surfaces to undergo rearrangement and removed adsorbed moieties. Effects of this treatment on metal powders have been investigated and evaluated by measuring the time before water droplets are absorbed on treated and untreated thin layers. Keywords: metal powder surface energy modification, additive manufacturing, high energy light irradiationIntroductionIndustrial processes involving metal part fabrication uses metal powders of a various range of particle size distributions, up to several hundred microns, such as injection molding where particles are compacted and subjected to heat treatments or additive manufacturing where the particles are spread into thin layers and selectively bonded in an iterative process.

Regardless of process, the particles’ surface properties, which can be summarily described as the powder surface energy, can impact the resulting part quality by affecting the optimal packing density or powder spreadability.Powder surface energy is influenced by several factors: particle size, morphology and struc- tural defects, and surface atomic arrangement. Micro-crystallographic structure within a particle can also impact surface energy [1, 2]. In addition, as powders are often kept in storage for ex- tended periods, interactions between the particles and the atmospheric environment as well as physiosorbed or chemisorbed impurities can alter the powder surface energy between manufacture and use [3, 4]. Some of these impurities may have a negligible or even positive affect on part quality, but often these impurities have a negative impact [5].

Some of these impurities may have a negligible or even positive affect on part quality, but often these impurities have a negative impact [5]. In additive manufacturing, powder recycling can accumulate additional residual impurities from the process; repeated recycling, while economical, can have cumulative detriment to powder characteristics and part quality.One avenue to consistent part quality is to ensure comparable or ideally, identical, powder surface energy, regardless of powder source and use. Consequently, there has been a great body of work focusing on thermal powder conditioning via prolonged bakes prior to use [6]. However, most of the work thus far are powder and application specific.

Consequently, there has been a great body of work focusing on thermal powder conditioning via prolonged bakes prior to use [6]. However, most of the work thus far are powder and application specific. In this paper, we focus on a more versatile ”equilibration” technique independent of metal, source, and application, as well as a demonstration of a fast powder surface evaluation methodology.Experimental ProcedureA range of metal powders were used in this work, although the primary focus was on gas atomized spherical powders with known size distributions.

In some cases, the powder, as a result1590Figure 1: SEM Images of some powders examined: (a) copper, (b) Hastelloy, (c) stainless steel 316L, and (d) Inconel.Composition Powder Stainless steel 316L SS1 Stainless steel 316L SS2 Stainless steel 316L SS3 Stainless steel 316L SS4 Stainless steel 17-4PH SS5 Ti-6Al-4V Ti1 Ti-6Al-4V Ti2 Ti Ti3 Mo Mo AlSi12 Al Inconel 718 Inconel Hastelloy C762 Hastelloy Cu; atomization process A Cu1 Cu2 Cu; atomization process B FeNiCoMo Fe:Ni:Co:Mo (68:16:8.5:4.5, Ti balance)Particle size D50 (µm) 17 21 34 32 32 38 36 27 38 46 32 32 14 14 33Table 1: Subset of commercial powders used in the experiment and their D50 particle sizes. Details of additives and contaminants are not available.1591of the atomization process or electrostatic forces, has small particle agglomerates on the main par- ticles, as seen in Figure 1a and 1d. Adding chemical agents can reduce the electrostatic forces and may be present in some of the samples [7, 8].

Adding chemical agents can reduce the electrostatic forces and may be present in some of the samples [7, 8]. Table 1 lists the samples used along with their D50 powder size.A Novacentrix PulseForge 1300, a commercial Xe flash system, was selected as the energy source for this experiment because of its ability to deliver rapid, high-energy optical pulses with a broadband spectrum spanning from 250-1000 nm. The particles’ surface energy can be altered by exposure to high temperatures to remove impurities or modify surface atomic arrangement. However, it is important to avoid excessive temperatures to prevent powder sintering or melting. Consequently, maintaining the particle structural integrity while altering its surface energy requires an instantaneous rise in surface temperature between 40-80% of bulk metal melting point while the interior temperature remains low [9].

This can be achieved with short, high-energy pulses the Xe flash lamp provides and finite heat transfer velocity particularly when irradiating a thin layer of powder. This was confirmed by IR imaging where samples subjected to the most extreme parame- ters showed rapid dropping of surface temperature within 0.5 s of pulse termination did not exceed 110◦C, even after repeated pulsing.The energy dose delivered per 20 ms pulse was varied between 5-50 J/cm2. Considering that the pulse has a rise time of 0.5-1 ms and a fall time of 1-2 ms, the total steady-state duration ends up 7.5-15% shorter than the programmed 20 ms. The baseline number of flashes samples were subjected to was 5 flashes; however, effects of fewer flashes on powder treatment were also investi- gated. Delay between pulses ranged from 0.8-4 s depending on the pulse energy and is a limitation of the system recharge speed.

Delay between pulses ranged from 0.8-4 s depending on the pulse energy and is a limitation of the system recharge speed. Three to five samples were measured for each experimental param- eter.The PulseForge 1300 is equipped to accommodate different atmospheric environments. As such, both an inert (Ar) and reductive (Ar-H2) environment were explored at atmospheric pressure. Samples were not irradiated when exposed to air ambient in order to prevent sample oxidation. The lack of oxide formation was visually confirmed via lack of surface discoloration and with EDX. Both ambient and heated samples were investigated, the latter taken up to 100◦C with the built-in resistive heater.Metal powders were mechanically spread into thin layers between 70-800 µm on a glass sub- strate, then placed into the PulseForge system where a rough pump-purge sequence, down to sub- 10 millitorr, is initiated before filling with the desired atmosphere and irradiated.

To determine the effects the irradiation had on the powder surface energy, the samples were removed from the Xe flash system and evaluated with a Kr¨uss Drop Shape Analyzer (DSA100) within 30 min of treat- ment [10]. A 4 µL water droplet is automatically deposited by the DSA100 onto the surface and its change in contact angle over time monitored with the DSA100’s camera system. Because the water surface tension is a constant between samples, the length of time for a droplet to be absorbed is a proxy for the powder surface energy, referred to as the ”droplet residence time”. Although this is a qualitative approximation technique, the droplet residence time is sufficient to demonstrate powder surface energy changes as a function of Xe pulse irradiation.

Although this is a qualitative approximation technique, the droplet residence time is sufficient to demonstrate powder surface energy changes as a function of Xe pulse irradiation. Water was selected for its high surface tension to facilitate the evaluation of high and low surface energy powders.1592ResultsThe results of the powder-water interaction can be categorized into three cases, as shown in Figure 2 below. First, when the powder surface energy is comparable to that of water’s surface tension, the droplet residence time is low, typically <0.5 s, showing that the droplet is quickly absorbed (Fig. 2a). In contrast, when the powder surface energy is much lower than water surface tension, the converse is true: the droplet residence time is high, exceeding tens of seconds, and the droplet remains unchanged on the surface (Fig. 2b). In the third case, the droplet disrupted the powder surface upon contact.

2b). In the third case, the droplet disrupted the powder surface upon contact. This phenomenon is likely due to shallow penetration of irradiation, resulting in a thin layer of treated powder with high surface energy (HSE) over the remaining untreated, low surface energy (LSE) powder. As the droplet wets the top HSE region and reaches the interface between the two contrasting surface energies, the interaction with the LSE powder deforms the surface (Fig.

As the droplet wets the top HSE region and reaches the interface between the two contrasting surface energies, the interaction with the LSE powder deforms the surface (Fig. 2c).Figure 2: Video stills and schematics illustrating the categories of droplet behaviors 0.2 s after droplet deposition showing the case of (a) high surface energy where the water droplet is quickly absorbed into the powder layer, (b) low surface energy where the water droplet remains on the surface for an extended period, and (c) both high and low surface energy within the sample layer, causing surface disruption when the water droplet reaches the interface.Ray tracing calculations showed that the pulse irradiation can reach 80-170 µm below the surface before full absorption. The absorption depth is dependent on variables like the metal, re- flectivity, particle size, morphology, and packing density.

The absorption depth is dependent on variables like the metal, re- flectivity, particle size, morphology, and packing density. Additionally, heat transfer from convec- tion and conduction between particles may increase the depth of heated particles to 100-250 µm. These calculations were confirmed by monitoring the depth of partial fusion of samples exposed to increased flash energies and further evidenced by samples that fall under the third category.Prior to irradiation, the powders used has a wide range of initial surface energies; some al- ready have short droplet residence times while most have severe surface energy mismatch with1593water. The following experimental parameters were initially held constant: the flash envelope con- ditions (20 ms), sample atmosphere (inert; Ar), powder temperature (100◦C), and the number of pulses (5x). Energy density per pulse was varied from 9-20 J/cm2.

Energy density per pulse was varied from 9-20 J/cm2. Figure 3 below shows the direct correlation between the energy density and droplet residence time. There is a clear indicator of a required threshold energy for the powder surface energy to change; this threshold energy is ma- terial dependent. At higher energy densities, the residence times are comparable across powders, indicating it is feasible to achieve similar final surface energies despite their initial states.Figure 3: Droplet residence time as a function of energy density/pulse for a subset of treated powders. As-is refers to the powder without any flash treating.Considering the energy cost and efficiency, the powder temperature and number of flashes were reduced to see if equilibration of powder surface energies still occurred. Pulse energy density was held constant at 20 J/cm2 because of its consistent effect on surface energy independent of material, as were the other variables.

Pulse energy density was held constant at 20 J/cm2 because of its consistent effect on surface energy independent of material, as were the other variables. Decreasing the powder temperature from 100◦C to ambient (∼ 25◦C) had largely minimal effects, as long as the threshold energy was met. Figure 4a shows that at low energy conditions (1x; RT), droplet residence times and droplet behaviors degrade. Reduction of flashes from 5x to 3x and 1x saw minimal changes, though the sensitivity to flash reduction was also material dependent (Fig. 4b). Further tuning of flash parameters may yield optimized results balancing efficiency and cost.As foreign impurities are a likely contributor to initial low surface energies, a reductive en- vironment with the same initial experimental parameters (100◦C, 5x flash) was briefly explored to see if the forming gas could further react with impurities and improve results.

However, changes to the experimental atmosphere had no perceivable difference in results. It is possible that the changes are difficult to capture with this qualitative methodology due to the complexity of metal particle surfaces and impurities interactions. In addition, it is possible that the impact of atmo- spheric changes is metal-dependent.1594Figure 4: (a) Droplet residence times increase as the total energy decreases. Broadly speaking, the treated powders still outperformed nontreated powders, but a threshold energy is evident for consistent results. (b) For the same material (stainless steel), the number of flashes imparted on the sample has negligible effect on droplet residence time.To further investigate the effect of impurities on metal powders, the specific case of carbon contamination in titanium alloys was chosen because of its use case in additive manufacturing [11, 12].

Impurities like carbon, especially as residue from binder jet agents, pose a challenge for tita- nium additive manufacturing due to its impact on mechanical properties. Two titanium alloys, one being a virgin, high-performing powder for printing (Ti-1) and one having a known carbon con- tamination due to recycling (Ti-2), were both flash treated and demonstrated an increase in surface energy (Fig. 3). Furthermore, Ti-2 was intentionally doped with additional carbon using 0.01 wt% sodium dodecyl sulfate and then irradiated. Chemical compositions of all samples were analyzed with elemental combustion technique (LECO) and showed reduction of carbon content after Xe treatment, tabulated in Table 2.

Chemical compositions of all samples were analyzed with elemental combustion technique (LECO) and showed reduction of carbon content after Xe treatment, tabulated in Table 2. While this can only remove some surface contaminates, repeated flashing may further reduce the carbon content.Carbon wt% (LECO) Sample 0.008 Ti1 – as-is 0.007 Ti1 – flashed once 0.016 Ti2 – as-is 0.013 Ti2 – flashed once 0.010 Ti2 – flashed three times Ti2 – intentionally contaminated 0.038 Ti2 – intentionally contaminated and flashed 3X 0.027Table 2: Combustion analysis (LECO) results of virgin, recycled, and carbon-doped titanium pow- ders before (as-is) and after flash treatment.1595Aside from removing absorbed moieties, in- creasing powder surface temperatures to 50-70◦C of the bulk melting point can trigger rearrangement of surface atomic crystalline structures to a more ther- modynamic configuration [13].

Because the rear- rangement occurs at high temperatures, over time, the effects of the flash treatment may deteriorate as the surface atoms spontaneously reconfigure to thermodynamic equilibrium under ambient condi- tions. Selected treated powders stored in lab ambi- ent conditions were periodically measured over the course of a few weeks. Figure 5 on the right shows the gradual decrease in surface energy of the pow- ders. Rate of surface energy change is likely depen- dent on the initial powder conditions, metal type, and flash parameters.Figure 5: Droplet residence time as a func- tion of aging samples under ambient condi- tions show the time-sensitivity of the flash treatment.ConclusionWe have demonstrated with a high powered Xe flash lamp the feasibility of increasing the surface energy of a thin metal powder layer to a consistent level regardless of the metal used and powder history.

The changes in surface energy are likely due to rearrangement of atomic structure of particle crystalline surfaces and removal or adsorbed impurities. Critically, the samples did not sinter or melt during the treatment, due to the localized nature of the short pulse durations and finite heat propagation velocity within the particle, thus ensuring that only the exposed particle surfaces and not the interiors were heated. This method can be implemented in an iterative process where untreated powder can be spread onto a treated layer for processing until enough material has been processed. The iterative layer-by-layer nature of this technique is highly compatible with that of binder jet printing and can be easily implemented where each powder layer is irradiated in the 3D printer before printing.1596References[1] L. Vitos, A. Ruban, H. Skriver, and J. Koll´ar. “The surface energy of metals”. In: Surface Science 411.1 (1998), pp. 186–202. DOI: https : / / doi . org / 10.

DOI: https://doi.org/10.1016/j.addlet.2022.100116.[13] E. Williams and N. Bartelt. “Thermodynamics of Surface Morphology”. In: Science 251 (1991), p. 393. DOI: https://doi.org/10.1126/science.251.4992.393.1597Supplemental MaterialIn addition to commonly used metals (stainless steel, titanium, and copper), a wide variety of other metals, both ferrous and non-ferrous, were investigated. In the figure below, it is clear that all materials treated demonstrate the same trend of reduced droplet residence time as energy density increases and that there is a clear material dependence on threshold energy for effective treatment.

The samples were subjected to the following flash parameters held constant while varying energy density: inert atmosphere (Ar), initial powder temperature of 100◦C, and five flashes.Figure 6: Xe flash heating results of several other metal powders.Two samples of stainless steel 316L and titanium powders (below) were further examined regard- ing the effects of powder temperature during flash heating and the number of flashes. Again, while the exact energy required for effective change in droplet residence time varied by material, once the threshold energy is reached, the initial powder temperature and the number of flashes do not affect the droplet residence time.

For example, powder temperature is seen to have a larger effect on the two stainless steel powders compared to the titanium powders before the threshold energy is reached.1598SS1; 100CSS1; RTSS2; RTDroplet Residence Time (s)Flash Number/Powder Temperatureas-is1X3X5X0.1110SS2; 100Cas-is1X3X5X0.1110Ti1; RTTi2; 100CTi1; 100CTi2; RTDroplet Residence Time (s)Flash Number/Powder TemperatureFigure 7: Energy density was held constant at 20 J/cm2 for all samples shown. Left: Droplet res- idence time dependence on initial powder temperature (room temperature or 100◦C) for stainless steel 316L. Right: Droplet residence time dependence on powder temperature for titanium.The figure above is a graphical representation of Table 2. As the powder is flash heated, the carbon contaminant content decreases, regardless of powder origin.1599.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEQUILIBRIUM SATURATION IN BINDER JETTING ADDITIVE MANUFACTURING PROCESSES: THEORETICAL MODEL VS. EXPERIMENTAL OBSERVEATIONSHadi Miyanaji*, Li Yang*Department of Industrial Engineering, University of Louisville, Louisville, KY40292AbstractIn Binder Jetting 3D Printing process, a feature is fabricated by depositing binder in the selective regions layer by layer until the geometry is completed. One of the main factors which influences the accuracy and strength of the green parts in this process is the spreading (i.e. lateral migration) and penetration (vertical migration) of the binder in powder bed, which is in turn determined by the capillary equilibrium between the saturated regions and the unsaturated powder.

lateral migration) and penetration (vertical migration) of the binder in powder bed, which is in turn determined by the capillary equilibrium between the saturated regions and the unsaturated powder. In the present study, a previously developed model for binder migration was applied to 420 stainless steel in order to evaluate the equilibrium saturation levels. Characterizations such as contact angle, specific surface area, succession pressure and capillary pressure were carried out in order to determine the theoretical equilibrium saturation amount. Furthermore, 10 single lines with dimensions of 46μm in diameter and 3mm in length were printed out and analyzed for empirical saturation level.

Furthermore, 10 single lines with dimensions of 46μm in diameter and 3mm in length were printed out and analyzed for empirical saturation level. The results of the theoretical model and experimental work were compared and discussed in details to validate its applicability in binder jetting process development.Keywords: Binder Jetting, 3D Printing, Equilibrium Saturation, Spreading, Penetration.IntroductionBinder Jetting Additive Manufacturing (BJ-AM) is an additive manufacturing processes in which a liquid binder is deposited through a print-head at the designated locations of the powder bed layer by layer for fabrication of part geometry. After deposition of binder for each layer, the surface of the powder bed is exposed to a fixed amount of heating which is applied via a heat lamp (e.g.

After deposition of binder for each layer, the surface of the powder bed is exposed to a fixed amount of heating which is applied via a heat lamp (e.g. infrared heater) so as to partially cure the binder and consequently establish appropriate mechanical strength to withstand the shear and gravitational compressive forces involved in spreading consecutive layers. This layer-by-layer process repeats until the part is completed.Based on the above description, geometry generation and integrity of the printed part in this process strongly depends upon the interaction between the liquid binder and powder. Once the binder droplets are placed on the selective regions, the liquid binder will migrate into the powder through pores located on the surface of the powder bed (so called imbibition) due to capillary attraction [1-6]. After imbibition is complete (i.e.

After imbibition is complete (i.e. after all the liquid binder permeates into porous medium), the drainage in which liquid binder migrates from the completely saturated region into the dry powder surrounding it, will begin to take place. Therefore, as binder migration progresses, the saturated regions drain and the dry regions imbibe until capillary pressure of both regions become equal. This is the state where equilibrium condition in which the capillary pressures are equal for both imbibition and drainage, is reached. The saturation level (the ratio of amount of liquid binder to the pore volume in the print material) in the equilibrium phase is of significant1945importance in BJ-AM process. This equilibrium saturation must be assumed as the saturation level determined by machine parameters for printing purpose.

This equilibrium saturation must be assumed as the saturation level determined by machine parameters for printing purpose. If the equilibrium saturation is much smaller than the saturation imposed by machine control parameters, the excess liquid binder will tend to migrate out of the designed boundaries of the part. On the other hand, If the equilibrium saturation is much greater than the machine-imposed saturation, the printed part won’t have sufficient mechanical strength due to lack of enough amount of binder, which will consequently result in weak bound between the particles and also successive layers. It has been experimentally shown that binder amount plays a major a role in controlling the part accuracy and strength [7-11]. Therefore, it is quite beneficial to be able to predict the equilibrium saturation for a given powder bed and liquid binder using physics-based modeling before the printing process.

Therefore, it is quite beneficial to be able to predict the equilibrium saturation for a given powder bed and liquid binder using physics-based modeling before the printing process. In the current study, a previously developed model is employed to predict the equilibrium saturation in 420 stainless steel powder bed and the results are compared with empirical measurements of saturation. One of the few attempts in this regard was reported by J. F. Bredt. In his PhD dissertation conducted at MIT, a model to predict the equilibrium fluid content of printed features from measurements of the capillary characteristics of the powder and physical properties of the fluid binder was developed [12]. It was proposed in this model that a balance of capillary pressure exists between a partially saturated region and the dry powder surrounding the feature.

It was proposed in this model that a balance of capillary pressure exists between a partially saturated region and the dry powder surrounding the feature. Following this preliminary work, in this paper experimental measurements of the capillary characteristics of the powder were used to calculate the predictions of the equilibrium saturation of the features in order to verify the equilibrium model.The ModelThe interaction between binder and powder in BJ-AM process is dominated by capillary, and to lesser degree, gravitational forces. The liquid binder in the powder bed distributes itself in the pores and contacts the powder surface with a local contact angle of 𝜃 determined by the surface energies of the solid, liquid, and vapor phase boundaries (see Figure 1). Although the meniscus formed in the vapor-liquid interface may assume a complex profile, it possesses a mean curvature and capillary pressure, defined uniquely everywhere in the binder.

Although the meniscus formed in the vapor-liquid interface may assume a complex profile, it possesses a mean curvature and capillary pressure, defined uniquely everywhere in the binder. On the basis of the Laplace equation, capillary pressure for a local curved interface (local curvature) within a pore is presented by [13]:∆𝑝 = 𝑝𝑐 = 𝛾(1 𝑟1+1 𝑟2) (1)where Δp is the capillary pressure across the fluid interface, γ is the surface tension is the r1 and r2 are the principal radii of curvature. For a capillary tube of radius r, the equation (1) becomes as follows [13]:𝑝𝑐 =2𝛾𝑐𝑜𝑠𝜃𝑟(2)1946Figure. 1 Local contact angle between liquid binder and a particle in the powder bedFor a real porous material where a liquid binder and powder interact, the outcome of the above equation would be a statistical average which is obtained over the void spaces in the vicinity of the considered pore [14].

From the equation, capillary pressure depends on the geometry of pore, contact angle 𝜃, and the degree of saturation (Sw). On the other hand, since the relationship of capillary pressure on saturation level (pc = pc (Sw)) cannot be analytically described due to the shape irregularity and complexity of pores in an actual porous material, laboratory experiments could be implemented to derive the relationship for any given porous media. A typical curve of capillary pressure as a function of saturation is shown in Figure 2.Figure. 2 Typical capillary pressure-wetting fluid saturation curves [12]From the figure, the curve follows two different paths for imbibition and drainage illustrating hysteresis. The hysteresis in the capillary pressure are known to be rendered by two mechanisms, contact angle (𝜃) and geometry of void space.

The hysteresis in the capillary pressure are known to be rendered by two mechanisms, contact angle (𝜃) and geometry of void space. While the former depends on the1947direction of the displacement, the latter arises from many bottlenecks of pores in the powder bed (ink-bottle effect) [14]. Due to the hysteresis in the capillary pressure curve, the saturated region and dry region the feature can exist at equilibrium with different fluid contents since each of these regions has different pressure characteristics.There are two important features in Figure 2. Irreducible wetting saturation, Swo, is a critical wetting fluid saturation at which the fluid displacement in the porous medium approaches zero. The capillary pressure at this saturation is called the dry suction pressure P0 which is a critical characteristic of a porous medium in determining the liquid binder migration during feature formation in BJ-AM process. The second point in Figure.

The second point in Figure. 2 labelled A on the right side of the graph is the minimum pressure (also called bubbling pressure, PB) required to initiate the displacement of a wetting fluid (e.g. liquid binder) in a fully saturated porous medium.In the model developed in [10], capillary pressure has been considered as the only driving force for binder migration, and equation (2) was used as the governing relationship between the capillary pressure and liquid binder saturation. In this model, pore curvature (K) is defined in such a way that term r is eliminated from the equation [12]:𝑝𝑐 =−2𝛾𝐿𝑉𝑐𝑜𝑠𝜃 𝑟= −𝐾 𝛾𝐿𝑉 𝑐𝑜𝑠𝜃 (3)Hysteresis in the capillary pressure between imbibition and drainage of a fluid results in two pore curvature for a single pore and for an assemblage of pores.

The critical nonwetting curvature, KNW, is defined by the bubble pressure, PB, and critical wetting curvature KW is defined by P0 as follows [12]:𝐾𝑊 =2 𝑅𝑊=−𝑃0 𝛾𝐿𝑉 𝑐𝑜𝑠𝜃𝑎𝑑𝑣(4)𝐾𝑁𝑊 =2 𝑅𝑁𝑊=−𝑃𝐵 𝛾𝐿𝑉 𝑐𝑜𝑠𝜃𝑟𝑒𝑐(5)where 𝜃𝑎𝑑𝑣 is advancing contact angle, and 𝜃𝑟𝑒𝑐 is the receding contact angle.The quantities 𝑅𝑊 and 𝑅𝑁𝑊 describe the pore geometry of the powder bed. There two radii are the dimensions of analogous cylindrical capillaries exerting the same capillary pressures as the dry powder and the saturated region, respectively. In Fredric’s model, a fully saturated feature is modelled as a cavity with two capillaries, one with radius 𝑅𝑊 that represents the path of dry pores that lead out into bulk powder bed, and the other one with a radius 𝑅𝑁𝑊 that represents the trail of the saturated pores that guide air into the feature [12].Figure. 3 The model for equilibrium of a liquid binder in a hydraulic feature [10]1948Liquid pressure can be transmitted across this feature.

3 The model for equilibrium of a liquid binder in a hydraulic feature [10]1948Liquid pressure can be transmitted across this feature. Besides fully saturated regions,pores that have been partially drained but still in hydraulic contact are included in this model.If the pressure within the larger capillary (RNW) is greater than that of smaller capillary in Figure 3, fluid will spontaneously migrate from left to right. In the current model, the excess pressure in the feature is expressed in terms of the differences in the pressure between two capillaries [12]:∆𝑃 = −𝑃0 + 𝑃𝐵 (6)In terms of the critical radii becomes∆𝑃 = 2𝛾𝐿𝑉[𝑐𝑜𝑠𝜃𝑎𝑑𝑣 𝑅𝑊−𝑐𝑜𝑠𝜃𝑟𝑒𝑐 𝑅𝑁𝑊] (7)Or alternatively,∆𝑃 = 𝛾𝐿𝑉[𝐾𝑊 𝑐𝑜𝑠𝜃𝑎𝑑𝑣 − 𝐾𝑁𝑊 𝑐𝑜𝑠𝜃𝑟𝑒𝑐] (8)This excess pressure is driving force for fluid migration in a porous medium. Liquid binder which is often employed in printing process is highly wetting to the powder (i.e. 𝑐𝑜𝑠𝜃𝑎𝑑𝑣 ≈ 𝑐𝑜𝑠𝜃𝑟𝑒𝑐 ≈ 1).

Liquid binder which is often employed in printing process is highly wetting to the powder (i.e. 𝑐𝑜𝑠𝜃𝑎𝑑𝑣 ≈ 𝑐𝑜𝑠𝜃𝑟𝑒𝑐 ≈ 1). Under these circumstances, the difference in critical pore curvatures (𝐾𝑊-𝐾𝑁𝑊) defines the driving force for binder permeation out of a fully saturated region. As a feature becomes less than fully saturated, the excess capillary pressure is proportional to the difference [𝐾𝑊 𝑐𝑜𝑠𝜃𝑎𝑑𝑣 − 𝐾𝑑𝑟𝑎𝑖𝑛𝑎𝑔𝑒 𝑐𝑜𝑠𝜃𝑟𝑒𝑐] where 𝐾𝑑𝑟𝑎𝑖𝑛𝑎𝑔𝑒 is dependent on the saturation. Therefore, the condition for hydrostatic equilibrium of a feature is𝐾𝑑𝑟𝑎𝑖𝑛𝑎𝑔𝑒(𝑆∗) = 𝐾𝑊 (9)where S* is the equilibrium saturation.Therefore, given the powder bed and liquid binder characteristic such as the suction pressure, the capillary pressure curves, surface tension, contact angle, etc. it is possible to estimate the equilibrium saturation of a printed feature using the relationship by Eq.(9).MethodologyIn this study, ExOne M-Lab machine was used for printing 420 Stainless steel samples.

As shown in figure 4, particle size of the used powder varies between 10μm and 40μm. The packing density of the powder was measured to be 55% once spread with the traverse rate of 3mm/sec. In order to measure the experimental equilibrium saturation, some cylindrical specimens were printed out and their dimensions were measured using an optical microscope. For all the samples, process parameters 100% saturation level, 70% curing power intensity, 45 sec curing time, and 3mm/sec spreading speed were employed.1949Figure. 4 The morphology of 420 Stainless steelIn order to determine the theoretical equilibrium saturation, the powder and liquid binder characteristics (e.g. liquid binder surface tension and viscosity, contact angle, specific surface area of the powder, and capillary pressure-saturation curve) need to be specified. The Cannon-Fenske Routine Viscometer, size 150, was employed to measure the viscosity of the liquid binder according to ASTM D445 and ISO 3104 [15].

The Cannon-Fenske Routine Viscometer, size 150, was employed to measure the viscosity of the liquid binder according to ASTM D445 and ISO 3104 [15]. For determining the surface tension of the binder, stalagmometric method was chosen. In this method, the total weight and number of drops of the liquid binder falling from a capillary glass tube (figure 5) were recorded, and the surface tension of the fluid were consequently calculated.Figure. 5 Volumetric pipet used for measuring the surface tension of the binderThe specific surface area of the powder was determined using BET (Brunauer, Emmett and Teller) method. The principle of this method is based upon physical adsorption of a gas on the surface of the solid and by calculating the amount of adsorbate gas corresponding to a monomolecular layer on the surface.

For this purpose, commercial TriStar 3000 device was utilized.Capillary Penetration Method is widely used for average contact angle determination applying the Washburn equation. In this method, the mass of liquid binder absorbed into the porous medium is recorded as a function of time. The contact angle can then be derived from the modified Washburn theory [1, 16, 17]:m2 =𝐶 𝜌2 𝛾𝑙𝑣 𝑐𝑜𝑠𝜃 𝜂𝑡 (10)1950where m is the mass increase of the column, t is time, 𝛾𝑙𝑣 is the binder surface tension, 𝜂 is the binder viscosity, 𝜌 is the binder density, and C is the capillary constant. The capillary constant C can be determined using the following relationships [16].𝐶 =𝑟 𝜋2 𝑅𝑘 24 𝜀2(11)where r is the pore radius, 𝑅𝑘 powder pack.4 is the radius of the capillary tube (figure 6), 𝜀 is the porosity of theThe effective radius, reff , (Eq.

12), can be replaced with r in the equation above [16]𝑟𝑒𝑓𝑓 =2(1−𝜙)𝜙𝜌𝐴(12)where 𝜙 is the volume fraction of solid in the tube, ρ is the density of the solid material, and A is the specific surface area per gram of solid (obtained from BET measurement).The apparatus for the Capillary Penetration experiment is shown in the Figure below. The powder particles were poured into a glass capillary (0.56 cm inner diameter and 7 cm long) which was supported at the bottom by Humboldt filter paper. The decrease in mass of the container (owing to the liquid binder permeation upward into the capillary) was recorded using an electronic balance as a function of time. The time t = 0 corresponds to the instant of capillary tube submersion in the liquid binder. After the binder mass was sucked into the capillary glass tube, packing density of the bed, and specific surface area of solid, average contact angle was then calculated.Figure 6.

After the binder mass was sucked into the capillary glass tube, packing density of the bed, and specific surface area of solid, average contact angle was then calculated.Figure 6. Apparatus used for Capillary Penetration experimentThe rate of imbibition measured can be also used to determine the dry capillary pressureaccording to Washburn equation:x2 =2𝐾𝑠𝑎𝑡 𝑃0 𝜀 𝜂𝑡 (13)1951where x is the distance of the penetration front from the contact point, 𝜀 is the porosity of the sample, 𝜂 is the viscosity of the liquid binder, Ksat is the hydraulic permeability, and P0 is the dry suction pressure.For measuring the hydraulic permeability, Ksat, in the above formulation a set up shown in figure 7 was used. The tube is secured in the bottom end and filled with a liquid on top of the sample.

The tube is secured in the bottom end and filled with a liquid on top of the sample. In this experiment, if the height is measured over time, an equation of the following form can be employed to determine the hydraulic permeability.h = h0 e-nkt (14)where k is the permeability, L is the sample thickness, A is the sample cross-section, r is the inside radius of the tube, and n is a geometrically determined constant which is defined asn = A/πr2L (15)where h0 corresponds to the initial height of binder at time 0.Figure 7. Schematic diagram of the set up for hydraulic permeability measurementLastly, filter paper method was used to obtain the capillary pressure-saturation curve in drainage [18]. This method was initially developed in the soil science for measuring soil suction. The filter paper method is cheap, simple, and capable of covering the entire range of suction pressures [19].

This method was initially developed in the soil science for measuring soil suction. The filter paper method is cheap, simple, and capable of covering the entire range of suction pressures [19]. The testing method is based the assumption that a filter paper will achieve equilibrium with a sample having a specific suction by either liquid or vapor moisture exchange between the sample and the filter paper. Once equilibrium is achieved, the water content of the filter paper will correspond to suction pressure through filter paper calibration curve shown in Figure 8. For the present study, samples of 0%, 10%, 30%, 45%, 60%, 80, 90%, and 100% saturation levels were prepared, and their corresponding capillary pressure were calculated using1952the water content of the filter papers after equilibration according to ASTM D5298-10. Thereafter, the capillary pressure-saturation curve was plotted with these data.

Thereafter, the capillary pressure-saturation curve was plotted with these data. It is of interest to note that capillary pressure corresponding 0% saturation level denotes the dry suction pressure described previously.Figure. 8 Calibration curves for two types of filter paper [18]Result and discussionThe results of filter paper method were converted to equivalent pore curvatures and displayed in Figure 9. This plot shows how the pore curvature defined previously varies with saturation amount introduced into the 420 stainless steel.Table 1 summarizes the results of different methods employed for characterizing the binder and powder interaction.

These outcomes were then used to calculated the dry suction pressure, P0, and the average contact angle between powder particles and the liquid binder permeating the powder bed.195310) 1 -98Packing density of 55%m μ (7e r u t e v r u c654e r o P3210020406080100Saturation amount (%)Figure. 9 The capillary pressure-saturation curve in drainageMethodFactor measuredResultsTable. 1 summary of the results The Cannon-Fenske Routine Viscometer Viscosity of the binder used (Pa.s) 0.00431BETSpecific surface area (m2/g) 0.03Stalagmometric method surface tension of the binder (N/m) 0.03From the capillary penetration measurements, it was found out that the penetration rate (imbibition rate) and the hydraulic permeability (Ksat) of the powder bed were 0.02 cm2/s and 1.3× 10−9 𝑐𝑚2 respectively. Therefore, according to the Washburn equation (Eq.

Therefore, according to the Washburn equation (Eq. 13) the dry suction pressure P0 was calculated as [1]:P0 =2.7×10−5 1.3×10−9 = 2 × 104𝑃𝑎 = 21𝑘𝑝𝑎𝑔𝑐𝑚3, and A= 0.03 m2/g, 𝜀 = 0.45 the effective radius, reff,Also, given 𝜙 = 0.55, 𝜌 = 7.74 and the capillary constant C were determined as:𝑟𝑒𝑓𝑓 =2(1−𝜙)𝜙𝜌𝐴=2(1−0.55)0.55×7.74×0.03× 10−6 = 6.8 × 10−6 𝑚𝐶 =6.8×10−6 𝜋2 (0.29)4(0.45)2 2= 4.8 × 10−16 𝑚5Similar to the original form of Washburn equation [1], which represents linear dependence of square of height versus time, the modified Washburn equation (Eq. 10) [16,17] also shows linear trend of binder mass square versus time as shown in Figure 10. This kind of phenomenon has been empirically observed by many researchers [17,20,21]. The dotted line in Figure 9 corresponds to the y2 = At trend-line, where A is a fitting parameter defined as A = 𝐶 𝜌2 𝛾𝑙𝑣 𝑐𝑜𝑠𝜃.

The dotted line in Figure 9 corresponds to the y2 = At trend-line, where A is a fitting parameter defined as A = 𝐶 𝜌2 𝛾𝑙𝑣 𝑐𝑜𝑠𝜃 . Given A=0.0017g2/sec according to Figure 9 and the capillary constant C = 4.8 × 10−16 𝑚5, the average contact angle between the liquid binder and particles in the powder bed can be calculated as:𝜂1954𝑐𝑜𝑠𝜃 =𝐴 𝜂 𝐶 𝜌2 𝛾𝑙𝑣=0.0017× .00431 4.8×10−16×1012× 0.03= 0.5 ⇒ 𝜃 = 60°This amount is well within the range reported in the literature for the contact angle of stainless materials with liquids similar to the used binder in the present study in terms of physical properties [22, 23].0.3) 2 g (0.252m d e r a u q S0.20.15m2 = 0.0017t R² = 0.97s s a M0.10.050050100150200Time (sec)Figure.

10 The binder mass absorbed versus timeFinally, the equilibrium saturation could be determined using Eq.(9) as follows:𝐾𝑑𝑟𝑎𝑖𝑛𝑎𝑔𝑒(𝑆∗) = 𝐾𝑊 =−𝑃0 𝛾𝐿𝑉 𝑐𝑜𝑠𝜃𝑎𝑑𝑣= 1.4 𝜇𝑚−1From the capillary pressure-saturation curve, this pore curvature which corresponds to the equilibrium saturation amount can be observed at a saturation level of approximately 85%. The saturation obtained represents the theoretical binder amount present in the equilibrium condition. Further saturation measurements are required be performed to evaluate the accuracy of the model. To determine the empirical equilibrium saturation, 10 single lines, with the diameter of 46 𝜇𝑚 and length of 3mm, were printed out with process settings of 100% saturation level, 70% power intensity, 45sec for curing time, and 2 mm/s as spread speed. The diameters of the lines were measured by optical microcopy.

The diameters of the lines were measured by optical microcopy. Figure 11 displays two examples of the printed features.Given the packing density of the powder bed, the binder density, the binder amount deposited from the machine settings, and having measured the dimensions of the printed features, the equilibrium saturation can be determined as the ratio of binder amount deposited to the pore volume present in the fabricated features. In the present study, ratio was measured to be 60% representing the equilibrium saturation in practice.1955Figure. 11 Two examples of printed features for empirical equilibrium saturation measurementsThe results reveals that the equilibrium saturation estimate from the capillary pressure measurements varies by 29% from the empirical saturation obtained from experimental measurements.

This discrepancy might be explained as follows.In the current model, capillary pressure/saturation curve in drainage is used to determine the final saturation amount using the dry suction pressure which will always lead to higher saturation level. In the equilibrium state there exists a trail of partially saturated pore with an irreducible wetting saturation level where the liquid binder ceases to migrate. The simplest case of this phenomenon is shown in Figure 12 with only two pores. In this picture, pore 1 drains and pore 2 imbibe until the capillary pressures between two pores become equal. The capillary pressure corresponding to pore 2 is referred to as irreducible wetting saturation.

The capillary pressure corresponding to pore 2 is referred to as irreducible wetting saturation. Due to the capillary pressure hysteresis between drainage and imbibition, if the capillary pressure characteristic curve in drainage is used for predicting the equilibrium saturation (which is the case in the model), this estimate will always be higher than actual saturation.The other factor which might be contributing to higher estimate of equilibrium saturation is neglecting gravity effect particularly in the vertical direction. The effect of gravity on the hydrostatic pressure within a feature can cause fluid to flow further downward out of the boundaries. Consideration of gravity effect in the model will result in higher driving pressures for binder migration which in turn lead to lower estimate of the saturation in the equilibrium condition.Figure.

Consideration of gravity effect in the model will result in higher driving pressures for binder migration which in turn lead to lower estimate of the saturation in the equilibrium condition.Figure. 12 Equilibrium state of two partially saturated pores1956ConclusionIn 3D printed features using Binder Jetting process, there exists an equilibrium condition in which the liquid binder ceased to migrate due to the balance between the capillary pressures of saturated regions and unsaturated ones. The saturation level in this state (also called equilibrium saturation) plays an important role in dimensional accuracy, integrity, and strength of the printed parts. In the present work, a previously developed model of binder migration was applied to 420 stainless steel in order to evaluate the equilibrium saturation levels.

In the present work, a previously developed model of binder migration was applied to 420 stainless steel in order to evaluate the equilibrium saturation levels. Different powder and binder characterizations such as contact angle, specific surface area, dry suction pressure and capillary pressure curve, were performed as to calculate the theoretical equilibrium saturation amount. Additionally, in order to the determine the equilibrium saturation in practice 10 single lines were printed out and analyzed using an optical microscope. The results showed that that the equilibrium saturation estimated from the capillary pressure measurements varies by 29% from the empirical saturation obtained from experimental measurements.

The results showed that that the equilibrium saturation estimated from the capillary pressure measurements varies by 29% from the empirical saturation obtained from experimental measurements. Using capillary pressure/saturation curve in drainage and also neglecting the gravity effect in the model might be two contributing factors to a higher estimate of equilibrium saturation.AcknowledgementThe authors also express their gratitude to the staff of Rapid Prototyping Center and Conn Center at the University of Louisville for their assistance and also Dr. Samuel Dilip for the help on performing this project.1957References[1]. E. W. Washburn (1921), The dynamics of capillary flow, Phys. Rev., 17: 273 – 283.[2]. A. MARMUR (1988), Penetration of a Small Drop into a Capillary, Journal of Colloid and Interface Science,Volume 122, Issue 1, Pages 209-219[3]. A.

ROR DETECTION IN LASER BEAM MELTING SYSTEMS BY HIGH RESOLUTION IMAGINGS. Kleszczynski*, J. zur Jacobsmühlen†, J. T. Sehrt* and G. Witt*Institute for Product Engineering, University of Duisburg-Essen, Germany †Institute of Imaging and Computer Vision, RWTH Aachen University, GermanyAbstractLaser Beam Melting as a member of Additive Manufacturing processes allows the fabrication of three-dimensional metallic parts with almost unlimited geometrical complexity and very good mechanical properties. However, its potential in areas of application such as aerospace or medicine has not yet been exploited due to the lack of process stability and quality management. For that reason samples with pre-defined process irregularities are built and the resulting errors are detected using high-resolution imaging. This paper presents an overview of typical process errors and proposes a catalog of measures to reduce process breakdowns.

This paper presents an overview of typical process errors and proposes a catalog of measures to reduce process breakdowns. Based on this systematical summary a future contribution to quality assurance and process documentation is aspired.1. IntroductionLayer-wise creation of solid bodies by joining formless material is the main feature of Additive Manufacturing (AM) technologies. Depending on the different kinds of processes the basic material used could be a powder, a viscous liquid or a solid material. The geometrical data of each layer is gained by slicing digital 3D-models in the suitable thickness. By means of the layer-wise production principle very complex parts can be produced, which may not be manufacturable with other processes. Another benefit of AM technologies is the direct and tool- free generation of parts from CAD data.

Another benefit of AM technologies is the direct and tool- free generation of parts from CAD data. Today these characteristics of AM technologies are used for many applications individual parts [1], [2], [3].including product development and production ofIn the 90ies most applications of AM technologies were found within the domain of Rapid Prototyping, which is until today often confused with AM technologies in general. Beneath this possible field of application the use of AM technologies for so called Rapid Manufacturing or Direct Digital Manufacturing increasingly gains importance [3]. With the development of production techniques for high density metal parts from one-component metallic powders in 1999 [4], a trend of producing individual metal parts for mass customization started. Beam Melting is a neutral term for this special AM technology and was established in the German VDI guideline 3404 [5].

Beam Melting is a neutral term for this special AM technology and was established in the German VDI guideline 3404 [5]. Other terms are Direct Metal Laser-Sintering (DMLS), Selective Laser Melting (SLM), LaserCUSING® or Electron Beam Melting (EBM) which are often related to a specific machine manufacturer. The beam source may be a laser or an electron beam. For this paper only Laser Beam Melting systems are taken into consideration. In the Laser Beam Melting process parts are created through a periodic sequence of powder deposition, layer creation and lowering of the building platform for the next powder deposition as shown in Fig. 1. Due to very high cooling rates during the melting and solidification, residual stresses induce curling effects of created layers. To avoid collisions of the recoating mechanism and these curled part areas so-called975support structures are necessary. The powder material used in the process can be reused for further processes.

The powder material used in the process can be reused for further processes. Sieving the used powder prevents large particles, welding spatters or other impurities from influencing the powder quality. For creation of layers a beam source melts the powder locally according to the layer specific scanning pattern. The compound of layers and beam paths occurs due to the solidification of the molten discrete areas. Typical applications for Laser Beam Melting can be found in the domain of medical implants, FEM optimized lightweight parts or tool inserts with contoured cooling channels [3]. In spite of its potential a breakthrough of the technology has not yet occurred in many fields of application. One of the main reasons is the lack of process stability and quality management. Especially in areas of high safety requirements like medical engineering and aerospace, customers and users of AM technologies need proof of the accuracy of produced parts.

Especially in areas of high safety requirements like medical engineering and aerospace, customers and users of AM technologies need proof of the accuracy of produced parts. To overcome this barrier it is necessary to find suitable answers to some key questions concerning process stability and quality assurance: which kinds of errors can appear in Laser Beam Melting? What is considered a critical process error? How can these critical errors be detected? And which measures can be taken for debugging?Fig. 1: Schematic representation of Laser Beam Melting.To find answers to these key questions, we provide a systematic overview of possible errors within the field of Laser Beam Melting. In a second step, errors from each category are detected in sample build processes using high resolution imaging. Upon these results possible measures for debugging during the Laser Beam Melting process are discussed.9762.

Upon these results possible measures for debugging during the Laser Beam Melting process are discussed.9762. State of the ArtLaser Welding as an established manufacturing process shows a few similarities to Laser Beam Melting. One thing both processes have in common is the importance of a defined heat input. For example, too much laser power will eventually cause hot cracks, too little energy might be responsible for poor composite weldseams. In the area of Laser Welding, these and other process errors are known and have already been categorized in international standards like DIN EN ISO 13919-1 [6].On the basis of this knowledge several approaches for optical process control systems were developed and commercialized. Most of them are based on cameras or photodiodes and use a beam splitter which enables a coaxial view through the laser beam.

Most of them are based on cameras or photodiodes and use a beam splitter which enables a coaxial view through the laser beam. By imaging the emitted light from the welding interaction region in this manner, closed-loop process control systems are able to regulate process characteristics such as laser power, welding speed or filler wire feed rate [7], [8].In the area of Laser Beam Melting these principles were adopted for the development of first approaches to process monitoring. Kruth et al. used a combination of a visual inspection CCD camera system and a melt pool monitoring system [9], [10]. The CCD camera system inspects the deposition of powder – due to wear and local damage the powder recoating system could cause irregularities in the powder surface deposited on the molten parts. These irregularities remain on the part surface after layer creation and induce high surface roughness.

These irregularities remain on the part surface after layer creation and induce high surface roughness. To detect this process error a light source has been mounted in front of the process chamber to illuminate the horizontal lines caused by the damaged powder recoating system. The perspective error caused by an angular camera position has been corrected by means of calibration algorithms. An analysis of grey values was used for the detection of powder bed irregularities. The melt pool monitoring system uses a CMOS camera and a photodiode. By means of a semi- reflective mirror the laser light is deflected towards a beam splitter, which separates the melt pool radiation towards the photodiode and the CMOS-camera. This setup enables coaxial image acquisition through the laser beam. Melt pool dimensions are measured by means of the CMOS camera. The photodiode detects the mean radiation, which is emitted from the melt pool.

Melt pool dimensions are measured by means of the CMOS camera. The photodiode detects the mean radiation, which is emitted from the melt pool. The possible resolving power reached by this setup is about 10 µm per pixel [10]. Similarly to Laser Welding it is possible to control process parameters like laser output power with this approach. This is useful for areas with varying thermal conductivity like overhanging structures. [11]Lott et al. expanded Kruth’s approach by imaging melt pool dynamics at higher scanning velocities by means of an additional laser illumination. The achieved resolution of this system was about 12 µm per pixel [12]. Another approach has been given by Pavlov et al. [13]. Here, a bi-color pyrometer system was installed coaxially to the laser beam.

Another approach has been given by Pavlov et al. [13]. Here, a bi-color pyrometer system was installed coaxially to the laser beam. For the analysis of thermal processes the parameters of hatch distance and powder thickness were varied.All mentioned approaches are using complex process monitoring systems which require a modification of the optical components of Laser Beam Melting systems. The complexity of these modifications presents a major drawback of these systems. Additionally, the field of view of all systems is limited to the current melt pool and its direct surroundings. The cooled down surface is not inspected, leaving the process result unsupervised. Performance of the described monitoring systems has been measured for selected process errors, only. A comprehensive documentation of process errors detected by the different approaches has not been given yet.

A comprehensive documentation of process errors detected by the different approaches has not been given yet. Furthermore some concepts are patented or licensed which makes it difficult to use them in different Laser Beam Melting systems.977For some types of process errors an in situ detection and documentation by means of a high resolution imaging system provides some advantages. We therefore present in this paper an alternative approach to error detection, which is system-independent and easy to implement.3. Experimental SetupAll studies described in this paper were made using an EOSINT M 270 Laser Beam Melting System from EOS GmbH. The powder material used was EOS NickelAlloy IN625. To provoke different kinds of process errors, parameters were varied in a wide field around the recommended standard. Process stability was investigated by building test bodies with critical geometrical features.

Process stability was investigated by building test bodies with critical geometrical features. Additionally, some types of errors were provoked by manipulating actors and optical components or using different powder qualities. For the latter waste powder was collected after sieving which contains larger particles.a)b)Fig. 2: a) Setup of the CCD camera system in front of machine window. b) Example image.For visual detection and analysis of errors a monochrome CCD camera system was used. The camera (SVCam-hr29050, SVS-VISTEK GmbH) was mounted in front of the machine window on the outside and captured images of the build platform from an observation angle (Fig. 2a). A tilt and shift lens (Hartblei Macro 4/120 TS Superrotator) helped to reduce perspective distortion by shifting the camera back and allowed placing the focal plane on the build platform without stopping down using its tilt ability. A 20 mm extension tube reduces the minimum object distance of the lens.

A 20 mm extension tube reduces the minimum object distance of the lens. The camera uses a 36 by 24 mm Kodak 29 megapixel sensor (6576 by 4384 pixels, pixel size 5.5 µm by 5.5 µm).We built all test bodies on a small building platform (100 mm by 100 mm) to evaluate the capabilities of our system at highest possible resolution. Later, the system can be extended to monitor the entire building platform (250 mm by 250 mm) at the cost of a reduction in spatial resolution.Resolving power of this setup was measured experimentally for a field of view of 130 mm by 114 mm with maximum lens tilt (8 °) and shift (10 mm) to minimize perspective distortion for a position similar to the final setup in front of the machine window. We used an USAF 1951 target and a microscope calibration slide with four circles on it (diameters: 1.5, 0.6, 0.15 and 0.07 mm) and acquired five images of each target by placing it in the center and at all978four corners of the field of view.

Perspective distortion was corrected by transforming the corners of the work piece carrier to a square in the image using four point homography estimation and warping with bicubic interpolation, resulting in an orthogonal view of the test target with a resolution of about 19 µm/pixel (see Fig. 3). The camera system is able to resolve two black lines on white background (40µm wide, 40 µm apart) from a distance of 50 cm. All circles on the microscope slide were detected and confirmed the pixel size. As can be seen from Fig. 2b this system enables us to inspect the part surface at bead level.a)b)Fig. 3: Example of perspective correction. (a) Building platform with corner markers and some powder in top right corner. (b) Corrected image of work piece carrier provides an orthogonal view of the build platform. Note: images intensities have been rescaled for visualization.Images presented in this paper were corrected using the same method.

Note: images intensities have been rescaled for visualization.Images presented in this paper were corrected using the same method. The edge length of the target image can be chosen arbitrarily; we used the minimum edge length of the distorted quadrilateral image to reduce the number of interpolated pixels, resulting in an image resolution of 4234 x 4234 pixels (~24 µm/pixel). The slightly larger pixel size in the object plane compared to the resolving power measurement is due to the larger field of view in the machine.For optimum image quality adequate lighting is required which must provide homogenous lighting for the mirror-like metallic weld bead structures to minimize specular reflections which would saturate the camera’s CCD sensor. It was found that diffuse lighting with a light source placed close to the working surface and opposite to the camera produces the best quality for surface images.

It was found that diffuse lighting with a light source placed close to the working surface and opposite to the camera produces the best quality for surface images. As the contrast of adjacent beads in the camera image depends highly on the shadow which is casted by each bead, the light source should always be perpendicular to the current bead orientation.In the machine, the powder blade movement prohibits the placement of any lighting component on the back of the machine. Therefore we realized diffuse lighting using matt reflectors on the back of the machine and powder blade (see Fig. 4) which are illuminated by two directed light sources from the front and right, respectively. Optimum lighting for alternating scan pattern is provided by two orthogonal light sources, one of which is selected depending on the current bead angle. Dark field illumination, which uses lighting parallel to the building plane to highlight surface disturbances, is applied separately.

Dark field illumination, which uses lighting parallel to the building plane to highlight surface disturbances, is applied separately. For each layer we acquire at least two images: one after powder deposition and one after laser beam melting. For lighting evaluation and better illustration we took images of typical errors with all possible lightings.979Fig. 4: Lighting setup for surface inspection and detection of elevated areas. Three light sources are used independently based on the current detection scheme. Contrast between adjacent beads is enhanced by illumination from the right side for bead angles from 45 ° to 135 ° and illumination from the front for angles from 0 ° to 45 ° and 135 ° to 180 °. Reflectors enable diffuse lighting on the work plane; dark field lighting uses parallel lighting to highlight embossments, it is used separately from the other two lighting sources.4.

Reflectors enable diffuse lighting on the work plane; dark field lighting uses parallel lighting to highlight embossments, it is used separately from the other two lighting sources.4. Results and DiscussionProcess Errors in Laser Beam MeltingBefore presenting images of typical process errors, the kinds of errors which can occur in the Laser Beam Melting process have to be defined. The following deliberations assume worst case scenarios in which errors are caused by human or technical failures. Basically the occurrence of these errors can affect process stability or part quality.Process stability is primarily endangered by collisions of the recoater mechanism and curled areas. As described in the first section of this paper, very high cooling rates during melting and solidification are causing residual stresses which induces plastic deformations within the created layers.

As described in the first section of this paper, very high cooling rates during melting and solidification are causing residual stresses which induces plastic deformations within the created layers. If the necessary support structures cannot compensate these deformations, a collision with the recoater mechanism stops the running process. Depending on the intensity of the occurred superelevation, support structures or parts are broken and lose their connection to the building platform. Critical geometric features can be another reason for local superelevations. Depending on the used material and process parameters there is a critical angle for overhanging structures up to which support structures are necessary. As a rough guide value 45° overhang can be considered as critical threshold. Especially for the creation of very complex parts a compromise between stable support connections and easy support removal has to be found.

Especially for the creation of very complex parts a compromise between stable support connections and easy support removal has to be found. Some geometrical features are even intended as functional areas within the created part and should therefore exhibit a certain surface quality which is influenced negatively by support connection. Beside the avoidance of curling effects, support structures are important for heat conduction, due to the fact that the conductivity of metal powders is about three orders of magnitude lower than that of solid metal [4]. If the heat induced by the laser beam is not conducted into the building platform, so called balling effects appear (Fig. 5a). Here, surface tensions of the viscous melt cause spherical areas within the part after solidification. The same problem occurs when too much heat in proportion to the surface area is conducted into the part. As for critical overhanging structures this heat accumulation induces balling effects.

The same problem occurs when too much heat in proportion to the surface area is conducted into the part. As for critical overhanging structures this heat accumulation induces balling effects. Some Laser Beam Melting Systems use980a stiff ceramic or metal blade for the recoating process. Due to the inflexible blade, elevated or curled part areas are causing a jam during the recoating process. Other Systems are using flexible silicon blades for recoating. With this setup jams can be mostly avoided, but there will be a higher wear of the flexible blade, which finally leads to uneven powder layers after powder recoating. Another source of error can be found in the connection of supports to the part. There are many support parameters which influence thickness, connection or removability, so that finding suitable parameters for support structures often depends on the specific experience of the technology users.

In consequence, the connection of supports to the built part can be referred to as a critical process phase, in which an insufficient compound could cause superelevations or a detachment of the built part. In both cases the process is demolished and has to be stopped. Furthermore process stability can be endangered by insufficient powder supply. This could be caused by a defective actuator in the powder reservoir platform mechanism or by powder reservoir exhaustion. Sensing devices usually notice when the powder stock is exhausted, but in worst case scenario a technical defect could prevent the running process from being stopped to refill the powder reservoir. Another reason for insufficient powder supply could be that the powder has not been compressed properly. Depending on the particle size distribution cavities could occur which collapse when the powder reservoir moves. As a consequence the powder level could not be sufficient for powder deposition.

Depending on the particle size distribution cavities could occur which collapse when the powder reservoir moves. As a consequence the powder level could not be sufficient for powder deposition. If there is no powder deposited on the created layers the laser beam will re-melt the solidified areas. Due to the repeated heat input by means of a defocused laser beam and the missing powder supply the so created parts cannot be completed.a)b)c)Fig. 5: a) Balling formation due to insufficient heat conduction caused by overhanging structures, b) good composite weldseams, c) poor composite weldseams by a wrong ratio of hatch distance h and width w combined with a high scanning velocity981Part quality is mainly measured by mechanical properties, surface roughness and dimensional accuracy. Depending on the utilized material and the heat input per second, hot cracks could eventually harm product quality.

Depending on the utilized material and the heat input per second, hot cracks could eventually harm product quality. The formation of cracks is mainly the result of high cooling rates during the Laser Beam Melting process. Brittle materials are particularly susceptible to this error type. Most materials available for Laser Beam Melting process show a good weldability and are assigned to ductile metal alloys. However, powder contaminations or different concentration of elements within the powder particles could cause these defects. To ensure proper mechanical properties of the created parts, the creation of good composite weldseams is essential. For this purpose a defined ratio between width and distance of melt traces has to be achieved (Fig. 5b). The width of melt traces depends on laser beam diameter, laser scanning velocity, powder layer thickness and laser power.

5b). The width of melt traces depends on laser beam diameter, laser scanning velocity, powder layer thickness and laser power. If the distance of melt traces – so called hatch distance – is too large, the compound of the respective melt traces will be deranged, which will finally cause porosities, resulting in poorer mechanical properties. A sudden decrease in laser power, e.g. caused by laser window pollution due to condensate deposits, a technical defect in the build platform actuator or an uneven platform resulting in a thicker powder layer at constant laser power leads to the same error type (Fig. 5c). Laser scanning velocity influences how much energy per second is conducted into the powder layer. A technical defect in the scanning system could therefore cause superelevations if scanning speed is too low, or poor composite weldseams if scanning speed is too high.Fig.

A technical defect in the scanning system could therefore cause superelevations if scanning speed is too low, or poor composite weldseams if scanning speed is too high.Fig. 6: Typical Process Errors in Laser Beam Melting categorized by influence, type and cause.Another main influence on product quality is the surface roughness. It is known that due to powder adhesion on part edges the measured surface roughness is larger in these areas. Better surface qualities can be achieved on surfaces parallel to the building platform. If there are disturbances in these layers, the surface quality could however be influenced negatively. Possible disturbances can occur due to a damaged recoater blade, caused by wear or superelevated areas, or contamination of the deposited powder. In case of a damaged recoater blade horizontal stripes are visible resulting in a locally thicker powder layer, which finally leads to unevenness in the982created layer.

In case of a damaged recoater blade horizontal stripes are visible resulting in a locally thicker powder layer, which finally leads to unevenness in the982created layer. A contamination of powder could be caused by non-metallic or coarse particles. Due to these foreign particles the necessary heat of fusion is locally different and therefore causes irregularities after solidification. Dimensional accuracy is another quality feature, which is influenced negatively by material shrinkage after solidification and incorrect geometrical slice data. Fig. 6 presents a systematic summary of all mentioned process errors.Error Detection by High Resolution ImagingAccording to the subdivision of process errors this section presents different types of errors in Laser Beam Melting using the described high resolution imaging system. In a first building test, process stability was examined by building different sample geometries with critical geometric features and critical support connection.

In a first building test, process stability was examined by building different sample geometries with critical geometric features and critical support connection. Additionally three cylindrical geometries were built for determination of the impact of contaminated powder.Fig. 7: Test samples built for examination of critical geometrical features and powder contaminationAs can be seen in Fig. 7 three test geometries have critical overhanging angles and were not connected to support structures. Part one and part two were placed parallel to the recoating mechanism which is unfavorable, due to the fact that in case of superelevation there is a large contact area between recoater and elevated part areas. This might cause a jam of the recoating mechanism. Part one furthermore features holes on the right side, which are critical when they are growing together. Part two features a critical overhanging area without support structures.

Part one furthermore features holes on the right side, which are critical when they are growing together. Part two features a critical overhanging area without support structures. Part three was inspired by turbine impellers, which feature overhanging blades. Support structures in these areas are difficult to remove. In consequence these parts are often built without any support structures which endangers process stability.Fig. 8a shows an image of this first test process taken in a z-built position of 5 mm after layer creation. Fig. 8b shows the same setup after powder recoating. As can be seen some areas of part one to three are superelevated and had contact to the recoating mechanism. Fig. 8c shows a detail of Fig. 8b. Depending on part geometry the superelevated areas will rise after each created layer. Furthermore there are horizontal lines visible in Fig. 8c which were caused by a damaged recoater blade. Part four in Fig.

Furthermore there are horizontal lines visible in Fig. 8c which were caused by a damaged recoater blade. Part four in Fig. 7 features supported overhanging surfaces in different983levels. The influence of different support connection parameters was investigated by this test sample. To achieve poor support the stability laser power used for support creation was decreased. After recoating some support particles were torn away by the recoater blade. These particles are also visible in Fig. 8d. The marked area in Fig. 8d shows the first part layer created on the poor support structures. Fig. 8e shows a magnified view of this area. Support connection was not stable enough to avoid curling effects of the created layer. With the next recoating procedure the recoater would tear off the created layer or even cause a jam. Finally Fig. 8f shows a magnified image of one cylindrical test sample after being exposed with contaminated powder. Comparing this image with Fig.

Finally Fig. 8f shows a magnified image of one cylindrical test sample after being exposed with contaminated powder. Comparing this image with Fig. 8e shows that the powder particles seem to be much darker and coarser than those of Fig. 8e (note that all images presented have been brightened for printing in the same manner). The solidified area in Fig. 8f shows a much more irregular surface. Melt traces are much more difficult to observe.a)b)c)d)e)f)Fig. 8: Various types of errors influencing process stability. See text for full description. All pictures were taken with the described high resolution imaging system. Note: image intensities have been rescaled for visualization purposes.In a second built process nine cylindrical test sample were built as shown in Fig. 9b. Herein the influence of energy input was examined. The three cylindrical test samples on the left side were assigned with process parameters that cause a low energy input.

9b. Herein the influence of energy input was examined. The three cylindrical test samples on the left side were assigned with process parameters that cause a low energy input. In detail the upper test sample was built using a laser power decreased by approximately 40 %. The middle test sample was built with a hatch distance increased by approximately 30 %. Finally the lower test sample was built using a scanning velocity increased by 40 %. According to this, the test samples on the right side were built using high energy input parameters. The three samples in the middle were984build using optimized parameters. After taking images as a reference, the laser beam diameter was enlarged for these samples. Through this energy input was decreased. Fig. 9a shows the test sample with enlarged hatch distance. The part exhibits a rough surface and irregularities. According to the remarks at the beginning of this section the image shows a poor compound of melt traces.

The part exhibits a rough surface and irregularities. According to the remarks at the beginning of this section the image shows a poor compound of melt traces. It is remarkable that the upper and the lower samples nearly show the same appearance. Fig. 9c shows a representative sample with high energy input. The sample shows a smooth surface with superelevated areas at the sample edges. The superelevation of the edge regions can be better seen after powder recoating in Fig. 9f. Finally Fig. 9e shows the surface of a cylindrical sample built with optimized parameters. Herein good composite weldseams are visible. No regions of superelevation can be recognized. After enlarging the laser beam diameter the same sample is again illustrated in Fig. 9d. As can be seen the surface is quite similar to those of Fig. 9a; it has become more rough and coarse particles are visible due to some reflections at the particle’s surface.

9d. As can be seen the surface is quite similar to those of Fig. 9a; it has become more rough and coarse particles are visible due to some reflections at the particle’s surface. The same error appearance would occur for increased powder layer thickness.a)b)c)e) Fig. 9: Various types of errors influencing product quality. See text for full description. All pictures were taken with the described high resolution imaging system. Note: image intensities have been rescaled for visualization purposes.d)f)5. ConclusionsWith this paper an alternative approach to error detection in Laser Beam Melting has been given. The used high resolution imaging system is easy to implement and therefore compatible to any existing Laser Beam Melting system. Furthermore the described high resolution imaging system is able to measure geometrical features and could therefore be used for the control of985dimensional accuracy.

Furthermore the described high resolution imaging system is able to measure geometrical features and could therefore be used for the control of985dimensional accuracy. Investigations for this option will be discussed in future publications. According to the systematical subdivision of process errors made in this paper and the documented error types it can be stated that superelevations and poor support connection are the most critical errors in the field of Laser Beam Melting. A possible measure for debugging this type of error could be a reduction of heat input in the affected areas. Hereby process heat can better be conducted through the solidified part areas and might therefore reduce heat accumulation. For this purpose coupling of the process monitoring system and the process control software is necessary. A low energy input could reduce part quality and is therefore undesirable in fields of applications with high safety requirements.

A low energy input could reduce part quality and is therefore undesirable in fields of applications with high safety requirements. Checking machine components and process parameters after the detection of low energy input errors can be stated as a necessary measure. If this error is detected in an early state, repeated exposure of the low energy input areas could be a measure for debugging this type of error. Powder contaminations and a damage of the recoating mechanism can be fixed easily after detection and are therefore regarded as less critical types of errors. The detailed effects of different types of errors on mechanical properties and semi-automated image analysis will be discussed in following publications.6. AcknowledgmentThe authors would like to thank the AiF association of industrial research unions “Otto von Guericke” e.V.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEstablishing Property-Performance Relationships through Efficient Thermal Simulation of the Laser-Powder Bed Fusion ProcessMohammad Masoomi, Basil Paudel, Nima Shamsaei, Scott M. Thompson*Department of Mechanical Engineering, Auburn University, Auburn, AL, 36849National Center for Additive Manufacturing Excellence (NCAME), Auburn University, AL, 3684929th Annual International Solid Freeform Fabrication Symposium - An Additive Manufacturing ConferenceAbstractIn order to learn how to modify additive manufacturing designs and processes to ensure lab-scale specimens and final components have similar properties, it is important that process-property relationships be established through thermal simulations.

In this study, two unique numerical methods for efficiently predicting the thermal history of additively manufactured parts via simulation are presented and validated. These numerical methods make use of an idealized, constant/uniform heat flux which is applied at each new layer and ‘bulk-layers’ which consist of several layers and allow the use of coarser meshes and longer time steps. To demonstrate and test the numerical methods, simulations are ran for the laser-powder bed fusion (L-PBF) of stainless steel (SS) 17-4 PH parts with different volumes.

Due to the high heat flux required for melting powder metal, L-PBF parts experience extremely localized temperature gradients and highly dynamic heating/cooling rates. The single (i.e. sensible) and multi-phase (i.e. latent) heat transfer rates that occur during L-PBF depend on process parameters, such as laser power, laser beam size, and traverse speed. Other design parameters, such as scan patterns, build orientation, powder size, and more, can affect the resultant thermal history of AM parts, as well [2].Since the thermal response of a part during its AM influences its microstructure, residual stress and defect formation, it is important to predict and model the inherent heat transfer as a* corresponding author: smthompson@auburn.edu1786function of part material type, geometry, location, orientation, operating environment, laser scan strategy, laser power, etc.

By understanding these process-property relationships, one gains more control over the AM part/process design for building higher quality parts that meet constraints. These process-property relationships enable end-users to reduce the number of trial-and-error experiments required to learn optimal process/design parameters for a given material.In-lieu of solving a complex set of governing, partial differential equations for the L-PBF problem, numerical simulation is employed, e.g. finite element methods (FEM) and/or computational fluid dynamics (CFD). The finite element method (FEM) may be used for numerically solving the discretized conservation equations via computational resources. The FEM discretizes the domain into thousands, or millions, of elements in which the conservation equations are applied iteratively using algebraic methods.

The FEM discretizes the domain into thousands, or millions, of elements in which the conservation equations are applied iteratively using algebraic methods. Modeling and simulation techniques will play a critical role in advancing AM by providing a means to quantify how process variables affect resultant component properties [3]. Through effective process modeling, the underlying transport physics can be better understood and final, AM part production and qualification can be accelerated.A major challenge in simulating the AM process is developing an efficient model that can produce information and knowledge at the overall scale of the desired part/component and build- process to inform engineering decisions [4,5].

Quantities of interest at the part scale include: distortions that can halt machine operation or place the completed part outside the desired geometric envelope; residual stresses that can create initial conditions detrimental to service-life concerns such as failure and fatigue; and local effective material properties, or at least indicators of where they might significantly deviate from the nominal properties expected from the process. There can be kilometers of laser travel length and thousands of powder layers in a single build, thus presenting a computational challenge. To be useful, part-scale simulations must be executable in a reasonable amount of time while retaining sufficient physical fidelity as to yield trustworthy results. An eventual goal is to have process models that are so efficient they could be included as part of an automated process optimization. Moran et al. [6] used a superposition-based approach to simplify and expedite the simulation process.

Moran et al. [6] used a superposition-based approach to simplify and expedite the simulation process. By modeling a single track and using data from that line and applying it as a line heat source for each track, the simulation time decreased by an order-of-a-magnitude.To expedite the simulation process and gain the ability to perform simulation on large- scale parts, new techniques are proposed herein. Bundles of layers are combined prior to simulating layer-wise AM in which an effective, uniform heat flux is employed. The goal is to decrease the total number of layers simulated for large-scale parts, thus reducing computational demand and time investment. The methods herein can aid research into AM process/design offsetting to ensure lab-scale specimens and real-life components have similar thermal histories.

The methods herein can aid research into AM process/design offsetting to ensure lab-scale specimens and real-life components have similar thermal histories. Numerical results are compared to experimental findings from the literature.Theory1787For modeling the L-PBF process, typically each layer is simulated separately along with a partition/island scheme and scan strategies. A typical part consists of hundreds of layers and thousands of meters of scan paths. In the proposed method, the heat flux applied at the top layer is assumed uniform. The time required to build a rectangular section is then calculated as:𝑚𝑚 × 𝑙𝑙 𝑤𝑤 × 𝑣𝑣where l is the length of the rectangular layer, m is the thickness of the part, w is the hatching space, and v is scanning speed. The constant heat flux applied at the top of the layer is calculated as:𝑡𝑡 =𝑃𝑃 𝑚𝑚 × 𝑙𝑙where P is laser power.

The constant heat flux applied at the top of the layer is calculated as:𝑡𝑡 =𝑃𝑃 𝑚𝑚 × 𝑙𝑙where P is laser power. Another method used to facilitate simulations is a method of combining multiple layers together to form a “bulk-layer” – in which the numbers of layers required to model the AM of a part is decreased significantly. Increasing the number of layers in each bulk layer will decrease computational time at the cost of accuracy.𝑞𝑞" =ValidationTetrahedral meshes were used for the powder bed and were uniformly dispersed, being ~0.05 mm3 in size (mesh lengths of ~ 0.5 mm). To ensure mesh independent results, simulations were repeated with mesh with lengths of 0.25 mm. No major changes were observed in the simulation results. The numerical model was built using COMSOL® Multiphysics 5.2. Simulations were conducted using the ‘Hopper’ cluster at the High Performance Computing Center at Auburn University. Each job utilized 80 processors and a total of 128 GB memory.

Simulations were conducted using the ‘Hopper’ cluster at the High Performance Computing Center at Auburn University. Each job utilized 80 processors and a total of 128 GB memory. Employed simulation results were validated using data from L-PBF experiments conducted by Dunbar et al. in which a local temperature response was recorded using a thermocouple at the bottom of the substrate [7]. In these experiments, two Ti–6Al–4V parts with geometries of 6.35 x 6.35 x 1.5 mm3 and 31.8 x 31.8 x 0.24 mm3 were fabricated. The time interval for both cases was 10 seconds. The substrate size for fabrication of the parts was 89 x 26 x 0.81 mm3 and 89 x 36.9 x 3.2 mm3, respectively. The process parameters for the L-PBF process are presented in Table 1.Table 1.

The substrate size for fabrication of the parts was 89 x 26 x 0.81 mm3 and 89 x 36.9 x 3.2 mm3, respectively. The process parameters for the L-PBF process are presented in Table 1.Table 1. Parameters used for simulating the L-PBF of Ti–6Al–4V parts.System Substrate material Powder description Powder layer thickness Laser spot diameter Laser power Scan speedEOS M280 Ti–6Al–4V Gas-atomized, air-dried 30 µm 100 µm 280 W 1200 mm/sTwo different simulations were conducted to compare the accuracy of each proposed modeling technique. First a constant heat flux was applied to each layer and the process was1788(1)(2)continued layer-by-layer until the whole part was simulated (S1). Second, a constant heat flux was applied at the surface of bulk-layers (S2). The temperatures predicted at the bottom of the substrate for S1-S2 are presented in Table 2. The simulation of a 31.8 x 31.8 x 0.24 mm3 part for S1-S2 took approximately 3 hours and 30 minutes, respectively.

The simulation of a 31.8 x 31.8 x 0.24 mm3 part for S1-S2 took approximately 3 hours and 30 minutes, respectively. The difference between temperature prediction and experimental results for fabrication of the small part was 20 ºC and 35 ºC for S1 and S2, respectively. For the larger part, the temperature difference between the experimental results and the S1 and S2 simulation results was 25 ºC and 35 ºC, respectively.Table 2. Comparison between experimental results from Dunbar et al. [7] and different simulations conducted of Ti–6Al–4V 6.35 x 6.35 x 1.5 mm3 part.No Experimental result [7] 1 2105 ºC 305 ºCSimulation #1 Simulation #2 125 ºC 330 ºC140 ºC 340 ºCIt may be observed in Fig. 1 that the substrate and powder bed absorb thermal energy from the fabricated parts. Additionally, it is observed from Fig. 2 that by fabricating a larger part and adding more power to the system, the effects of heat accumulation become more marked.

Additionally, it is observed from Fig. 2 that by fabricating a larger part and adding more power to the system, the effects of heat accumulation become more marked. Results suggest that modelling each layer separately via an ‘effective’/uniform laser heat flux (S1) is more accurate than a bulk layer approach (S2). However, when a constant heat flux is applied at the top of each layer the predicted temperature is higher than the measured temperature. This error could be explained by the fact that heat has less time to diffuse compared to when a local heat flux is modeled. Although individual layer modeling results in more accurate results, the traditional approach requires significantly longer computation times when compared to the proposed techniques.Figure 1. Temperature distribution of Ti–6Al–4V 6.35 x 6.35 x 1.5 mm3 part immediately after fabrication has finished.1789Figure 2.

Temperature distribution of Ti–6Al–4V 6.35 x 6.35 x 1.5 mm3 part immediately after fabrication has finished.1789Figure 2. Temperature distribution of Ti–6Al–4V 31.8 x 31.8 x 0.24 mm3 part immediately after fabrication has finished.ResultsThe validated numerical approach was used to investigate the temperature distribution along two different-geometry stainless steel (SS) 17-4 PH parts during L-PBF. One part was a cylinder with two cross-sections, i.e. a ‘dog bone’, as shown Fig 3 (a). The cylinder had a total length of 84 mm with the top and bottom portions having a diameter of 11 mm and the middle section having a length of 36 mm and a diameter of 7 mm. The other part was a square rod with a cross-section of 11 × 11 mm2 and thickness of 84 mm, as shown in Fig. 3 (b). Process parameters used for fabrication of these SS 17-4 PH parts are presented in Table 3. The time interval between fabrication of each layer was set at 10 seconds.Table 3.

3 (b). Process parameters used for fabrication of these SS 17-4 PH parts are presented in Table 3. The time interval between fabrication of each layer was set at 10 seconds.Table 3. Parameters used for simulating the L-PBF of a SS 17-4 PH part.System Substrate material Powder description Powder layer thickness Laser spot diameter Laser power Scan speed Hatching spaceEOS M280 SS 17-4 PH Gas-atomized, air-dried 40 µm 100 µm 220 W 755 mm/s 110 µm1790Figure 3. Schematic of L-PBF simulated SS 17-4 PH parts.The temperature distribution along both parts immediately after fabrication is presented in Fig. 4. It is observed that the temperature along newly-deposited layers is higher relatively to previously-deposited layers. Additionally, it is observed that the cylinder accumulated more heat relative to the square rod. Most of the thermal energy is diffused through fabricated parts.

Additionally, it is observed that the cylinder accumulated more heat relative to the square rod. Most of the thermal energy is diffused through fabricated parts. During fabrication of cylinder, the build area is smaller compared to the square rod, thus, heat accumulation becomes more important.As shown in Fig. 5, the maximum temperature is almost 100 ºC higher on the cylinder relative to the square rod immediately after their L-PBF. This elevated temperature would generally correspond to lower temperature gradients and cooling rates along top portions of the cylinder. It is important to notice that the effects of heat accumulation are more significant along newly-deposited layers and tends to fade away with previously deposited layers. Additionally, it is observed in Fig. 5 that when the build area during fabrication of the cylinder changes, the temperature distribution will significantly be affected. Smaller build area translates to higher heat accumulation.1791Figure 4.

Smaller build area translates to higher heat accumulation.1791Figure 4. Temperature distribution of (a) cylinder and (b) square rod immediately after L-PBF.Figure 5. Height-wise (Z) temperature distribution along centerline of (a) cylinder and (b) square rod immediately after their L-PBF.The different thermal responses shown in Figs. 4-5 indicate that, in general, different- sized/shaped specimens (e.g. cylinder and square rod) will have different microstructural and mechanical properties. In order to fabricate different-sized/shaped parts with similar mechanical properties, which would be important when transitioning from witness/test specimens to full-scale components, it is essential to fabricate parts with the same thermal response, i.e. local cooling rates, absolute temperatures, temperature gradients, etc. In order to achieve this, process parameters must be altered.

local cooling rates, absolute temperatures, temperature gradients, etc. In order to achieve this, process parameters must be altered. A simple parameter to control/alter, and one employed herein, is the time interval between layers of fabrication. To demonstrate this, the layer-to-layer time interval the cylinder part was increased from 10 seconds to 15 seconds. It is observed from Fig. 6 and Fig.17927 that, in this case, the thermal response of the cylinder and square rod become much closer to each other.Figure 6. Temperature distribution of fabricated (a) cylinder and (b) square rod after fabrication of the parts.Figure 7. Temperature distribution of (a) cylinder and (b) square rod along the main axis after fabrication of last layer.ConclusionsIn this study, two new techniques were investigated to more efficiently simulate the temperature distribution of part-scale specimens. The techniques were then used to simulate the1793thermal response of part-scale specimens.

The techniques were then used to simulate the1793thermal response of part-scale specimens. The unique temperature response of different- sized/shaped specimens was demonstrated, and its circumvention shown by increasing the layer- to-layer time interval of the smaller volume part. In general, such an approach can result in parts of almost similar temperature response independent of geometry or build orientation. Major findings of this study are summarized below:1. By applying constant and uniform heat flux, it is possible to accurately predict heat accumulation during fabrication of the parts efficiently.2. Bundling layers together prior to simulation will significantly decrease computation times, while only slightly increasing error.3. Parts with smaller build areas are more likely to experience effects of heat accumulation. 4.

Parts with smaller build areas are more likely to experience effects of heat accumulation. 4. Increasing the time interval will decrease effects of heat accumulation.There continues to be challenges in establishing process-structure-property-performance relationships for AM parts. One may adjust laser power and scan speed during fabrication to achieve target thermal response; however, changing these parameters may affect pore formation in the parts which in turn affects their mechanical response.AcknowledgmentThis work was partially supported by the National Science Foundation under Grant #1657195, and the National Aeronautics Space Administration (NASA) Grant #80MSFC17M0023.References[1] M. Masoomi, S.M. Thompson, N. Shamsaei, J. Pegues, Effects of shielding gas flow on heat transfer during laser-powder bed fusion of Ti-6Al-4V, Addit. Manuf. (2018).[2] M. Masoomi, S.M. Thompson, N.

tablishing the performance requirements for stab resistant Additive Manufactured Body Armour (AMBA)A. Johnson*, G. A. Bingham*, and C. E. Majewski†Loughborough Design School, Loughborough University, Loughborough, UK, LE11 3TU †Advanced Additive Manufacturing group (AdAM), Department of Mechanical Engineering, University of Sheffield, Sheffield, UK, S1 3JDAbstractBody armour is worn to lessen the likelihood of sustaining a life threatening injury. Such protective solutions are used every day by law enforcement officers around the world, with strict guidelines governing their design and testing.

These activities are monitored by government departments such as the Home Office Scientific Development Branch (HOSDB) within the United Kingdom (UK), and the National Institute of Justice (NIJ) within the United States.Despite providing protection against significant levels of impact energy, a number of historical issues continue to be present with modern fibre-based soft body armour – which once addressed may demonstrate an enhancement wearer operational performance.This paper therefore presents research highlighting such issues, and demonstrates how Additive Manufacturing (AM) technologies, particularly Laser Sintering (LS), could potentially be used to address such operational concerns whilst providing protection against a real-world threat.Results documented within this paper demonstrate that 5.6 mm thick planar samples, Laser Sintered from a 50/50 mix of virgin and recycled PA 2200 successfully achieved penetration resistance to the UK HOSDB KR1 impact energy of 24 joules.

These results therefore influenced the design, manufacture, and testing of a series of AM textile samples featuring an imbricated layout, which also demonstrated successful knife penetration resistance to the HOSDB KR1 level – thus developing stab resistant Additive Manufactured Body Armour (AMBA).IntroductionWithin the United Kingdom (UK) there are in excess of 144,000 police officers [1], and a further 765,000 sworn law enforcement officers within the United States (US) [2] – an occupation significantly at risk of encountering an assault within the workplace [3]. As part of their personal protective equipment, police officers are issued with body armour – the level of protection may vary depending on operational duties [4–6].

As part of their personal protective equipment, police officers are issued with body armour – the level of protection may vary depending on operational duties [4–6]. Body armour worn by these professionals can provide protection against a multitude of threats ranging from high-velocity ballistics to blunt-force and bladed artifacts, and are intended to lessen the likelihood of sustaining a life threatening injury as a result of a major incident [7,8]. Such armour can typically be placed into two categories [9–11]:Hard body armour – traditionally consisting of heavy ceramic plates encapsulated within a fabric carrier which fastens around its wearer, often used in military environments297Soft body armour - typically manufactured from aramid-based fibres such as Kevlar™, where flexible armour panels are inserted into a fabric carrier, and are commonly used by the police.

An example of which is shown in Figure 1.Figure 1 – Soft Body Armour Example: Stab-Resistant Duty Vest carrier (L) and aramid-based insert (R) Courtesy of Keltic Clothing [12]One of the major risks facing police officers is that of sustaining a sharp force injury as a result of a bladed threat [5,13]. Soft stab resistant body armour worn by these individuals must adhere to strict performance and test requirements. These are defined in a series of closely related standards from the following UK and US government departments:UK: Home Office Scientific Development Branch (HOSDB) - Publication 39/07 [14]  US: National Institute of Justice (NIJ) - Body Armour Standard 0110.01 [6]Both HOSDB and NIJ standards aim to consider the various factors which can affect the severity of a sharp force incident, by defining the use of a standardised test procedure, drop-test impact rig, and standardised engineered blades.

Additionally, these standards document the three levels of protection principle for the development of stab resistant body armour [14]:Knife Resistance Level One (KR1) – Armour for low risk patrolling environments which offer periods of maximum wearKnife Resistance Level Two (KR2) – General duty body armour providing a medium level of protectionKnife Resistance Level Three (KR3) – Body armour providing a high level of protection, typically worn for short periods in high-risk situationsA breakdown of these test energies and over-test conditions for each protection level,as defined by the HOSDB and NIJ, are documented within Table 1.298Table 1 - HOSDB and NIJ Stab Impact Energy Protection Levels [6,14]Energy LevelE1Strike Energy (J)24 +/- 0.5Velocity (ms-1)5.0 +/- 0.05Total Drop Vehicle Mass (kg) 1.9Maximum Penetration (mm) 7KR1E2 - over test36 +/- 0.66.2 +/- 0.051.920E133 +/- 0.65.9 +/- 0.051.97KR2E2 - over test50 +/- 0.77.3 +/- 0.051.920E143 +/- 0.66.7 +/- 0.051.97KR3E2 - over test65 +/- 0.88.3 +/- 0.051.920Despite significant advances in technical fibre development and the implementation of HOSDB and NIJ standards, a number of historical issues continue to exist with many of the current protective solution.

Such issues include:Restrictive and cumbersome use – impairing physical mobility [15,16]  Poor fitting resulting in nerve and musculoskeletal injuries [15,16] Increased thermo-physiological loading, thus resulting in illness and reduced operational performance [17]One area yet to be fully explored in an attempt to reduce such a burden on the wearer is that of Additive Manufacturing (AM). The unrivalled design freedom and ability to realise truly novel and highly complex functional assemblies via AM, specifically Laser Sintering (LS), may provide an opportunity to develop personalised Additive Manufactured Body Armour (AMBA) [18,19].

The utilisation of this technology may facilitate the incorporation of operational performance enhancing attributes which [16]:Maximise impact energy dissipation and freedom of movement  Minimise impact deformation and penetration  Provide enhanced fit and wearer comfortHowever, to be considered as a viable solution, there is a need to assess the technical performance of AM body armour against existing body armour standards – beginning with the HOSDB and NIJ knife resistance standards previously documented.Experimental MethodologyFour experiments were performed using a guide-rail drop test impact rig to the HOSDB KR1 impact energy of 24 joules (J). At this energy level, the maximum permissible penetration protruding through the underside of a test sample is 7.0 mm. All experiments used HOSDB specification „P1B‟ engineered knives. To support the underside of each sample during testing, non-hardening, oil-based „Roma Plastilina® No. 1‟ clay was used as backing material.

To support the underside of each sample during testing, non-hardening, oil-based „Roma Plastilina® No. 1‟ clay was used as backing material. The Plastilina backing material was packed into a 150 x 150 x 80 mm tray, open on its top face, and conditioned for three hours at a temperature of 35°C [8].All test samples were manufactured from either 100% virgin or a 50/50 mix of virgin and recycled polyamide (PA 2200) supplied by EOS GmbH, and Laser Sintered using an EOS P100 Formiga system set to the following settings [8]:Layer thickness – 0.1 mm299Build chamber temperature – 176.5°C  Removal chamber temperature – 150°C  Scan strategy – mechanic  Contour laser power – 16W  Contour scan power – 1,500 mms-1  Hatching laser power – 21W  Hatching scan speed – 2,500 mms-1Two types of samples were tested across the four experiments. Experiments One to Three featured planar samples which were manufactured to the dimension of 80 x 80 mm.

Experiments One to Three featured planar samples which were manufactured to the dimension of 80 x 80 mm. These samples were impacted tested to a centrally located strike zone identified on their top faces, as shown in Figure 2.Figure 2 - Planar test sample highlighting the central impact zoneThese samples ranged in thickness from 1-10 mm, with three samples tested at eachthickness. An overview of these experiments is shown in Table 2.Table 2 - Overview of Experiments One to ThreeSample AttributeExperiment OneExperiment TwoExperiment ThreeDesign PowderThickness RangeThickness Increment Total Number of SamplesPlanar100% virgin1-10 mm1.0 mm30Planar50/50 virgin/recycled1-10 mm1.0 mm30Planar50/50 virgin/recycled 4.1-5.9 mm 5.1-5.9 mm 0.1 mm54The results generated from the first three experiments were used to guide the generation of a series of novel articulated samples (AM textiles) which featured a scale link design within Experiment Four.

These samples were manufactured via LS from a 50/50 mix of virgin and recycled PA 2200 to a dimension of 120 x 120 mm, and featured an individual scale thickness of 4.0 mm. Two strike zones were identified - directly on the top face of a centrally located scale, and in between two centrally located scales – as shown in Figure 3.300Figure 3 - Articulated test sample strike locations (L) and section view showing minimum material thickness (R)The articulated samples featured a minimum material thickness of 5.6 mm in its thinnest region – as influenced by the planar sample test results from Experiments One - Three. The total height of the articulated samples when manufactured was 16 mm.Experimental Results & DiscussionExperiment OneTest samples ranged from 1-10 mm thick in increments of 1.0 mm.

The total height of the articulated samples when manufactured was 16 mm.Experimental Results & DiscussionExperiment OneTest samples ranged from 1-10 mm thick in increments of 1.0 mm. This experiment demonstrated that when manufacturing samples from virgin PA 2200, a minimum sample thickness of 8.0 mm was required to consistently achieve an acceptable level of penetration resistance to the KR1 impact energy. Also noteworthy from this experiment is that 66% of the samples tested fractured into multiple pieces - as shown in Figure 4.Figure 4 - Experiment One 4.3 mm (L) and 7.0 mm (R) thick virgin power shattered samplesIn fact, 20 of the 21 samples tested prior to the 8.0 mm thick planar samples failed in this manner, thus prompting a further avenue of exploration – the use of a recycled and virgin mix of PA 2200.Experiment TwoSamples were manufactured in thicknesses ranging from 1-10 mm in 1.0 mm increments using a 50/50 mix of virgin and recycled PA 2200.

Impact testing demonstrated that a minimum sample thickness of 6.0 mm was required to consistently achieve penetration resistance below the 7.0 mm HOSDB KR1 permissible limit – as depicted in Figure 5.301Figure 5 - Experiment Two recycled powder sample test resultsInconsistent penetration resistance within acceptable limits was also documented for 5.0 mm thick samples. This therefore prompted the need for further investigation of the impact penetration resistance of samples around 5-6 mm thick.Experiment ThreeWithin this experiment samples manufactured from a 50/50 mix of virgin and recycled PA 2200 were tested. These ranged in thickness from 4.1-4.9 mm and 5.1-5.9 mm, increasing in 0.1 mm increments.

These ranged in thickness from 4.1-4.9 mm and 5.1-5.9 mm, increasing in 0.1 mm increments. The results from this experiment demonstrated that the 5.6 mm thick sample group were the first to consistently achieve penetration resistance within HOSDB and NIJ requirements for the KR1 impact energy of 24 J – as shown in Figure 6.Figure 6 - Experiment Three Stab Test Results (L) and 5.6 mm thick planar sample (R)Based on previous literature, the enhanced performance of 50/50 virgin and recycled PA 2200 mix samples may be attributed to an increase in the molecular weight due to the use of recycled powder. This powder had previously experienced temperatures in excess of302160ºC, causing the molecules to rearrange and pack more efficiently - thus increasing the tensile strength of the sintered parts [20].

Further investigation is required in order to verify whether this is the cause of enhanced penetration resistance.Experiment FourResults from the previous three experiments were used to drive the generation of a series of articulated AM textile samples, featuring a scale link based imbricated layout - manufactured from a 50/50 mix of virgin and recycled PA 2200. Samples were tested at both previously identified strike zones, with results demonstrating that all samples achieved penetration resistance within the acceptable limits as defined by the HOSDB and NIJ KR1 impact energy of 24 J. Impact testing of these samples is shown in Figure 7.Figure 7 - Impacted articulated sample (L), maximum penetration demonstrated (M), and link failure (R)Impacts directly on the top face of the scale links demonstrated no penetration, whilst a maximum knife penetration of 1.6 mm was experienced when the samples were stuck between a pair of scale links.

This result was experienced on one test sample, and was likely due to a failure in the linkage mechanism. Due to the overlapping nature of the articulated samples, knife penetration was restricted and any failure within the design was contained within its structure.ConclusionsInitial research into the feasibility of generating stab resistant AM textiles has successfully established a number of initial criteria that can be used to achieve stab resistance to the HOSDB KR1 impact energy of 24 J.

Results from initial planar sample testing were used to design, manufacture, and test a number of articulated samples which demonstrated stab resistance significantly below the HOSDB and NIJ 7.0 mm permissible limit.In addition to utilising AM technology to establish stab resistant requirements, the enhanced design and manufacturing freedom of AM solutions such as LS may now be used to begin investigating whether historical issues associated with wearing body armour could potentially be addressed.

For example, using LS to develop AMBA solutions which reduce thermo-physiological loading on the wearer, or to generate individualised protective solutions.To build on the initial results presented, research is now focused on establishing aseries of additional criteria to enhance the performance of AMBA.303Further WorkFurther work is being performed as part of PhD research within Loughborough University to investigate methods of enhancing the protective performance of stab resistant AMBA. A series of planar samples manufactured from PrimePart DC powder are currently being tested to the KR1 impact energy of 24 J to determine their knife resistance performance. Additionally, samples manufactured from the 50/50 mix of virgin and recycled PA 2200 are also being tested to the KR1 over-test impact energy of 36 J – where a maximum 20 mm penetration is permissible.

It is anticipated that the results from this over- test condition will provide an indication as to the requirements to attain protection against the enhanced KR2 impact energy of 33 J, where the maximum permissible penetration is 7 mm - thus establishing a medium level of protection as defined by both the HOSDB and NIJ.Research is also focused on using graphical generative algorithm solutions for the design and development of scale links that incorporate features to enhance strike energy dissipation and assembly flexibility, whilst minimising scale geometry – a number of examples of which are demonstrated in Figure 8.Figure 8 – Examples of scale design featuresFeatures currently being explored in order to further enhance the protective performanceof stab-resistant AMBA include: Disturbing strike incidence through the incorporation of disruptive design features  Enhancing the linkage mechanism between scales to encourage and maximise impact energy dissipationInvestigating the cross-sectional design of the scale linkAcknowledgementsInitial results were generated as part of a MSc.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceESTIMATING STRENGTH OF LATTICE STRUCTURE USING MATERIAL EXTRUSION BASED ON DEPOSITION MODELING AND FRACTURE MECHANICSSang-in Park *, †, Narumi Watanabe **, David W. Rosen *, †G.W.W. School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA ** The Boeing Company, Seattle, WA †Singapore University of Design and Technology, Singapore AbstractGeometrical complexity in lattice structures yields large bounding surfaces to be approximated during additive manufacturing (AM) processes. In material extrusion, approximation of geometries using finite-sized thin filaments introduces defects such as voids and gaps in as-fabricated geometries. This initiates cracks between layers and increases possibility of fracture by crack propagation.

This initiates cracks between layers and increases possibility of fracture by crack propagation. As a result, a lattice structure fabricated by material extrusion tends to fail at significantly lower stress than estimated strength without consideration of fracture mechanism. The goal of this research is to estimate strength of material extruded lattice structures considering bonding strength among layers. To achieve this, the bonding strength is determined based on a deposition process modeling scheme and fracture mechanics analysis. A two-layer deposition model is generated to investigate deposited geometry, and the effective interlayer-bonding strength is calculated using a cohesive zone model (CZM) and peel tests. The resulting strength is incorporated into the property-estimation procedure.Introductionto design and manufacture geometrically complex parts such as lattice structures and cellular materials.

The resulting strength is incorporated into the property-estimation procedure.Introductionto design and manufacture geometrically complex parts such as lattice structures and cellular materials. Recent advances in additive manufacturing process have been improving geometrical accuracy and mechanical properties of as- manufactured parts. However, geometrical approximation occurred in additive manufacturing processes still yields geometrical degradation [1, 2]. In particular, the geometrical degradation becomes more critical in lightweight lattice structures composed of many thin elements like bars and beams, since the structures have large bounding surfaces, which increase possibility of geometrical degradation [3]. In material extrusion processes, to approximate geometries in a building plane. Since the filaments have finite size and elliptical or circular cross sections, voids and gaps shown in Figure 1 are formed between filaments during manufacturing.

Since the filaments have finite size and elliptical or circular cross sections, voids and gaps shown in Figure 1 are formed between filaments during manufacturing. The process-induced voids and gaps initiate cracks in fabricated parts in the building plane as well as along the building direction. These cracks limit mechanical strength of fabricated parts as they grow along filament interfaces or layer surfaces. Furthermore, voids and gaps may have larger influence on strength of lattice structure than bulky parts, since lattice structures are composed of thin members of which the size is several times of the filament diameter. Thus, crack propagation more easily degrades strength of lattice structures than bulky parts.Additive manufacturing process provides new opportunitiesthin filaments are deposited along deposition pathsMany researchers have investigated degradation in mechanical properties of additively manufactured parts. Ahn et al.

Ahn et al. reported material anisotropy depending on deposition patterns in material extrusion processes [4]. In the research, significant strength drop was observed when a raster angle is perpendicular to an external force direction. Rodriguez et al. investigated a relationship between a raster angle of deposition path and tensile strength [5]. The authors addressed that the raster angle has a large influence on strength, and they proposed a design framework for improving mechanical properties of material extruded parts by optimizing raster angles. Bagsik and Schӧppner measured tensile strength of additively manufactured specimens, which are built in three building directions [6].

Bagsik and Schӧppner measured tensile strength of additively manufactured specimens, which are built in three building directions [6]. This research also reported large reduction in strength over 40% for vertically built specimens.2127(a) Stair steps [2](b) Voids and gapsFigure 1 Geometrical degradation in material extrusion processesResearchers have implemented fracture mechanics to understand the material anisotropy and strength reduction shown in additively manufactured parts. Thomas and Rodriguez investigated wetting and thermally driven diffusion bonding processes to estimates fracture toughness between two deposition lines [7]. The authors performed two dimensional heat transfer analysis to capture temperature distribution during solidification process in material extrusion process and analytically derived fracture toughness between deposition lines. Fracture mechanics approach recently receives more attentions. Young et al.

Fracture mechanics approach recently receives more attentions. Young et al. proposed an experimental procedure to measure interlayer fracture toughness in additively manufactured parts [8]. The authors modified a standard double cantilever beam (DCB) specimens in order to apply to additively manufactured specimens. Gardan et al. developed a design framework to improve fracture toughness based on the direction of principle stresses and demonstrated more durable deposition path pattern against crack propagation [9]. Kishore et al. devised an infrared preheating system to enhance the interlayer fracture strength of parts manufactured in a material extrusion based big area additive manufacturing (BAAM) machine [10]. Seppala et al. investigated interlayer bonding strength based on isothermal healing time in polymer-polymer welding process [11].

Seppala et al. investigated interlayer bonding strength based on isothermal healing time in polymer-polymer welding process [11]. Research related to fracture mechanics in additive manufacturing research area has focused on experimental procedure.Cohesive zone model (CZM) is a numerical fracture analysis model devised to describe crack propagation and fracture in bonded interfaces [12-14]. Several researchers recently have implemented the numerical scheme to study interlayer fracture and to investigate strength of additively manufactured parts. Liravi et al. developed a cohesive zone model to estimate a pulling-up force at failure in a bottom-up projection- based photo polymerization process [15]. Spackman et al. constructed CZM and proposed model-based design framework for a fiber-reinforced soft composite additive manufacturing process. [16] The authors selected CZM parameters based on calibration experiments. Ahmadi et al.

[16] The authors selected CZM parameters based on calibration experiments. Ahmadi et al. implemented CZM to assess mechanical properties of metal parts fabricated by a powder bed fusion process [17]. The authors utilized CZM to define interactions among melt pool boundaries based on their observation that melt pool boundaries are weaker than grain boundaries so that defects are initiated in melt pool boundaries. They predicted effects of microstructure on mechanical properties of fabricated parts based on numerical analysis with CZM.In material extrusion processes, fracture mechanics approach can be used to explain strength degradation between layers or deposition paths. However, previous research has focused on experimental aspects of the approach. Constructing a numerical fracture mechanics model can help to expedite design processes utilizing material extrusion processes by providing a systematic analysis framework.

Constructing a numerical fracture mechanics model can help to expedite design processes utilizing material extrusion processes by providing a systematic analysis framework. The goal of this research is to estimate failure strength of a lattice structure considering structural mechanics as well as fracture mechanics. To achieve the goal, we develop numerical fracture mechanic models based on the cohesive zone model. The numerical models are calibrated based on peel tests using double cantilever beam specimens. The strength degradation in material extrusion processes is assessed based on fracture toughness determined from numerical models and embedded crack length predicted from process simulation models.

The strength degradation in material extrusion processes is assessed based on fracture toughness determined from numerical models and embedded crack length predicted from process simulation models. Figure 2 shows a conceptual diagram for the research framework.2128Fracture energy & Fracture toughnessProcess SimulationNumerical model for deposition shapeEmbedded crack length characterizationCohesive Zone ModelNumerical model for crack propagationParametric study for finding cohesive zone parametersCalculate stress intensity factorbased on the crack lengthStrength estimationCalculate stress intensity factorbased on the crack lengthPeel TestTest for crack propagationObtain the opening force –displacement relationshipCalculate stress intensity factorbased on the crack lengthFigure 2 Conceptual diagram for research frameworkIn following sections, we explain an experimental procedure for a peel test, and describe a numerical modeling procedure using cohesive elements.

Next, we present parametric study to calibrate cohesive parameters in the numerical model based on test results. Then, we explain a process simulation model to determine initial crack length. Finally, we present the estimation procedure and the results.Peel TestIn this research, ASTM D5528-13 testing method is implemented to quantify mode I inter-laminar fracture toughness [18]. Figure 3 (a) shows a standard double cantilever beam specimen. The standard DCB specimen is devised to assess bonding strength of a continuous fiber-reinforced composite material. However, this DCB specimen is not suitable to measure bonding strength between deposition paths in material extrusion for two reasons. The first is that infill raster pattern is not controllable and the second is that bonding strength between two deposition paths is of interest in this work. To solve the problem, the standard specimen is modified as shown in Figure 3 (b).

To solve the problem, the standard specimen is modified as shown in Figure 3 (b). The modified DCB specimen has a small interlayer bonding area, which is filled with only two deposition paths so that it is guaranteed that only two deposition paths are bonded between layers.Initial crack embeddedTwo deposition linesin the center region(a) Standard DCB specimen(b) Modified DCB specimenFigure 3 Standard and modified peel test specimens2129Tape to embed initial crack (23 mm)Figure 4 Fabricated DCB specimenThe specimens were fabricated in Dimension 1200se from Stratasys® . The raw material was ABS P400 and a T16 nozzle tip was used, which leads 0.254mm layer thickness. To embed an initial crack in the specimens, we paused fabrication process after deposition for the 7th layer, and we applied a strap of a 3M® blue paper masking tape to locate an initial crack. The crack length is 23 mm. After that, we resumed the process to complete fabrication.

The crack length is 23 mm. After that, we resumed the process to complete fabrication. Figure 4 shows a fabricated DCB specimen. Peel tests are performed in an Instron® universal tensile test machine. The specimens were installed using metal wires as shown in Figure 7 (a).Cohesive Zone ModelIn this work, we utilize a cohesive zone model to simulate interlayer bonding fracture in DCB specimens. CZM is capable to describe a gradual separation phenomenon caused by crack growth. A traction-separation law (TSL), which defines a relationship between cracking opening and required traction characterizes this model. Several traction-separation laws have been proposed based on material characteristics. In this research, a bilinear traction-separation law shown in Figure 5 is used. In the bilinear traction-separation law, it is assumed that the relationship is composed of two linear regions. To define the two linear regions, three important parameters are required.

In the bilinear traction-separation law, it is assumed that the relationship is composed of two linear regions. To define the two linear regions, three important parameters are required. The parameters are penalty stiffness, cohesive strength and fracture energy.The penalty stiffness is related to an initial mechanical response of a cohesive zone before cohesive traction reaches cohesive strength. The magnitude of the penalty stiffness must be high enough to avoid interpenetration of crack surfaces [19]. In this work, the penalty stiffness is selected based on an approach proposed by Turon et al. [20] as following:EKtEwhere, a parameter much larger than 1.

In this work, the penalty stiffness is selected based on an approach proposed by Turon et al. [20] as following:EKtEwhere, a parameter much larger than 1. In this work,is an elastic modulus of a raw material,tis a thickness of plies adjacent to a bonding region, andis set to 50, which is recommended by Turon et al.Crack opening (m) Traction required (Pa)Penalty stiffness (K) Cohesive strength (σi)Area: fracture energy (Gc)Figure 5 Bilinear traction-separation law2130(1)isThe cohesive strength is related to damage initiation. After cohesive traction reaches the cohesive strength, required cohesive traction starts to be reduced due to damage accumulation. There are several criteria based on stress, strain or displacement field in deformed configuration.

There are several criteria based on stress, strain or displacement field in deformed configuration. In this work, the maximum stress criterion is used to determine damage initiation and the value of cohesive strength is set to is yield strength of the raw material.0.55Y, where The area below a traction-separation law is a fracture energy that is directly related to fracture toughness. Once the fracture energy is determined, damage propagation response can be expressed. In this paper, we determine a value that minimizes difference in the maximum traction force between peel test and finite element analysis with CZM for the fracture energy.Parametric Study for Calibrating Fracture AnalysisThe cohesive zone model explained in the previous section is implemented to a numerical model of a DCB specimen in ABAQUS® . The numerical DCB specimen model is shown in Figure 6.

The numerical DCB specimen model is shown in Figure 6. A thickness of cohesive zone is set to 1 , and the zone is modeled using 8-node three-dimensional cohesive elements (COH3D8). The nonlinear analysis is performed using the model and results are compared with test rsults. Figure 7 shows a deformed shape of DCB specimens in the test and analysis.mTo determine the fracture energy, we construct a response surface model based on results from four different fracture energy levels, which are 100, 200, 300, 400 N/m. The traction force – opening displacement responses from analysis are compared with those from peel tests in Figure 8. The smaller fracture energy is applied to the cohesive zone model, the weaker traction force is yielded in specimens. The fracture energy and fracture toughness of each specimens are determined from the response surface model as listed in Table 1.

[21]. Two dimensional process simulation models for material extrusion are developed using ANSYS® Polyflow analysis package. This process simulation model is capable to describe deposition shape and temperature distribution during material extrusion. The simulation modeling is composed of two stages. The first simulation model is deposition and cooling model of the first layer of filament.The first layer is deposited onto a build platform, which is assumed to be at a constant temperature of 100 °C. The geometry and mesh before the deposition is shown in Figure 9 (a). A volumetric flow rate at the nozzle,Qis calculated using Eqn. (2):rQvWH2132Mean206.220.334(2)rvWHwhere is its height, assuming that the deposited filament had more of a rectangular shape than circular [22].

(2):rQvWH2132Mean206.220.334(2)rvWHwhere is its height, assuming that the deposited filament had more of a rectangular shape than circular [22]. By applying the calculated volumetric flow rate at the nozzle entrance and gravitational force, and using the remeshing technique in ANSYS® Polyflow, the deposition of the first layer was performed.is the deposition velocity,is the width of the deposited road, andThe second simulation model was the deposition of the second layer of filament on top of the first layer. Once, the temperature distribution after the first layer cooling was exported from the previous simulation, and the geometry and mesh before the second layer deposition is shown in Figure 9 (b). The procedure for this simulation model was similar to that for the first layer deposition. However, it was crucial to simulate the conduction heat transfer between the two layers in this stage.

The procedure for this simulation model was similar to that for the first layer deposition. However, it was crucial to simulate the conduction heat transfer between the two layers in this stage. This was accomplished using the fluid-to-fluid contact capability in ANSYS® Polyflow.

This was accomplished using the fluid-to-fluid contact capability in ANSYS® Polyflow. Required material properties of ABS P400 are listed in Table 2.Layer Height = 0.254 mmBuild Platform (T = 100 °C, v = 0 mm/s)Tambient= 100 °CNozzleExtrude(a) 1st layer deposition(b) 2nd layer depositionFigure 9 Material extrusion process simulation modelingTable 2 Material properties of ABSViscosity Expression𝟏 𝜼 = 𝒆[𝟓𝟔𝟕𝟓.𝟐( 𝑻−𝟏 𝟓𝟎𝟑.𝟏𝟓)]𝟐𝟐𝟐𝟖𝟗(𝜸̇ )−𝟎.𝟔𝟖Coefficient of Thermal Expression7.38 x 10-5 m/(m-°C)Thermal Conductivity0.17 W/(m-°C)Specific Heat1423.51 J/(kg-°C)Density1040 kg/m3Glass Transition Temperature (Tg)105.0 °C2133Deposited cross sectional shape from process simulation modelInitial crack lengtha(a) Deposited shape and temperature distribution(b) Embeded crack lengthFigure 10 Resulting shape and temperature distributionFigure 10 (a) shows resulting deposition shape and temperature distribution.

The crack length between layers is measured as shown in Figure 10 (b). The measured initial crack length is 0.08532 mm. This crack length is incorporated into fracture failure criteria in the following section.Estimation of Strength of Lattice StructureIn this section, a failure criterion of lattice structures considering fracture is developed and presented. Mechanical properties of lattice structures depend on their microscopic characteristics. Once a strut in a lattice structure starts to fail, the lattice structure also starts to fail. Two failure modes are considered in this research. The first is elastic failure and the other is fracture failure.

Two failure modes are considered in this research. The first is elastic failure and the other is fracture failure. The conceptual description of failure criteria is presented in Figure 11.YieldEquivalent tensile force Component in build directionBuilding anglemaxYImaxICKKFractureFor detecting fractureFor detecting elastic failureBuilding directionStrut directionFigure 11 Failure criteria of a strut in a lattice structure2134In this work, the elastic failure criterion is developed based on yielding of a strut. For detecting elastic failure, resultant axial stress by external forces is compared with yield strength of a raw material, which is 22 MPa for ABS P400. It is noteworthy that the total axial force is considered for elastic failure. The fracture failure criterion is established by utilizing fracture toughness of a cracked cylinder. A strut in a lattice structure is considered as a cracked cylinder.

The fracture failure criterion is established by utilizing fracture toughness of a cracked cylinder. A strut in a lattice structure is considered as a cracked cylinder. It is assumed that the cracked cylinder is fractured when resultant stress intensity factor reaches the fracture toughness. The stress intensity factor is calculated based on Gao and Herrmann’s work [23] as following:244maxmax41,124IccrIraKIIIIrwhere,maxIKis the maximum resultant stress intensity factor,maxis maximum resultant axial stress,is the embedded crack length. To calculate the stress intensity factor of a strut using Eqn. radius of strut, and (3), the crack length is required. In this work, the crack length determined in the previous section is applied. For detecting fracture failure, only force component in the building direction is applied to calculate stress intensity factor.

For detecting fracture failure, only force component in the building direction is applied to calculate stress intensity factor. The fracture toughness for this criterion is set to calculated fracture toughness with the CZM model in the previous section.aFigure 12 shows a tensile specimen for validating the proposed estimation approach. The specimen has 2D rectangular lattice structures in the center. Tensile specimens are fabricated in Dimension® 1200es with the same process parameters for DCB specimens at five different building angles: 0◦, 30◦, 45◦, 60◦, 90◦. The fabricated specimens are tested in the Instron universal tensile test machine.Figure 13 presents comparison among estimation and experimental results. The experimental result shows significant drop in strength when the specimens are built at high building angles. However, elastic failure criterion is not able to capture this strength degradation since the main failure mode of specimens built at high angle is fracture.

However, elastic failure criterion is not able to capture this strength degradation since the main failure mode of specimens built at high angle is fracture. The fracture criterion improves estimation by describing fracture mode failure.Build angleFigure 12 Tensile specimen with 2D rectangular lattice structure2135r(3)is the00.20.40.60.811.2030456090Normalize StrengthBuilding angle (degreee)Estimated strengthTestFracture failureElastic failureFigure 13 Comparison of estimated strength of 2D rectangular lattice structureEstimation of Strength of Lattice Structure using Cohesive Zone ModelIn this section, strength of tensile test specimens with two dimensional rectangular lattice structures is estimated using cohesive zone model developed in previous sections. Numerical models of the specimens at 0◦ and 90◦ building angles are modeled using an as-fabricated voxel modeling approach as shown in Figure 14. Using symmetricity, half of a rectangle is generated.

Numerical models of the specimens at 0◦ and 90◦ building angles are modeled using an as-fabricated voxel modeling approach as shown in Figure 14. Using symmetricity, half of a rectangle is generated. A layer of cohesive elements is inserted between layers near a joint. The obtained cohesive parameters in the previous section are implemented to the cohesive elements.Comparison of normalized strength among estimates and measurements are listed in Table 3. At the zero building angle specimen which is built horizontally parallel to building plane, estimates using Eqn. (3) is similar to the cohesive zone model as shown less than 10% error. However, at the 90◦ degree building angle, estimates using Eqn. (3) shows large error over 130% larger than the experiment measurement.Layer of cohesive elementsBuild angleSymmetryFigure 14 As-fabricated voxel model with cohesive elements2136Table 3 Comparison of normalized strengthBuilding angle0◦90◦Estimated strength using Eqn.

(3)0.953(-4.69%)0.406(135.56%)Cohesive zone model0.914(-8.58%)0.160(-7.36%)Experiment1.0000.172ConclusionThis research proposes a strength-estimation approach for a lattice structure considering interlayer fracture. We developed a numerical model for a double cantilever beam specimen with cohesive elements and determined the cohesive parameters by comparing mechanical responses with experimental results. The embedded crack length between layers was determined by process simulation model. The estimates for strength of a lattice structure were calculated analytically and numerically. For analytic approach, the stress intensity factor of a strut with calculated crack length was compared with fracture toughness from numerical analysis of double cantilever beam models. For numerical approach, strength of a lattice structure was determined from the maximum force of an as-fabricated voxel model with cohesive elements.

For numerical approach, strength of a lattice structure was determined from the maximum force of an as-fabricated voxel model with cohesive elements. Comparison found that significant strength degradation at high building angles can be predicted by considering interlayer fracture phenomenon, and that the proposed approach using the cohesive zone model is capable of capturing interlayer fracture mode and improving accuracy of the strength-estimation procedure. In the future research, we will investigate more lattice structures using the proposed approach.AcknowledgementThe authors gratefully acknowledge support from the National Science Foundation, grants CMMI- 1200758 and CMMI-1538744.

Alfano, M., Furgiuele, F., Leonardi, A., Maletta, C., and Paulino, G.H. Cohesive zone modeling of mode I fracture in adhesive bonded joints. in Key Engineering Materials. 2007. Trans Tech Publ. Campilho, R.D.S.G., de Moura, M.F.S.F., and Domingues, J.J.M.S., Using a cohesive damage model to predict the tensile behaviour of CFRP single-strap repairs. International Journal of Solids and Structures, 2008. 45(5): p. 1497-1512. Park, K., Choi, H., and Paulino, G.H., Assessment of cohesive traction-separation relationships in ABAQUS: A comparative study. Mechanics Research Communications, 2016. 78, Part B: p. 71-78. Liravi, F., Das, S., and Zhou, C., Separation force analysis and prediction based on cohesive element model for constrained-surface Stereolithography processes. Computer-Aided Design, 2015. 69: p. 134- 142. Spackman, C.C., Nowak, J.F., Mills, K., and Samuel, J., A cohesive zone model for the stamping process encountered during 3d printing of fiber-reinforced soft composites.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEstimating the Effects of Part Size on Direct Laser Deposition Parameter Selection via a Thermal Resistance Network ApproachGarrett J. Marshall1, Scott M. Thompson2, †, Steve R. Daniewicz3, Nima Shamsaei21Center for Advanced Vehicular Systems (CAVS), Mississippi State University, MS, 397622Department of Mechanical Engineering, Auburn University, Auburn, AL, 368493Department of Mechanical Engineering, University of Alabama, Tuscaloosa, AL, 35487† Corresponding author: Email: scott.thompson@auburn.edu Phone: (334) 844-4820AbstractA mathematical model for heat transfer during the Directed Energy Deposition (DED) process is proposed.

The model employs the thermal resistance network analogy and is developed to aid one in predicting part size effects on its temperature distribution during manufacture, and in how to compensate such effects via suitable process parameter selection. The model predicts a pseudo steady-state temperature response in the melt pool. The temperature variation along the heat affected zone of a thin-walled part is estimated while assuming deposition is occurring far from the substrate. Predicted melt pool and bulk part temperatures are validated against Laser Engineering Net Shaping (LENSTM) experimental data obtained via a dual-wavelength pyrometer and in-chamber infrared camera, respectively. Results demonstrate that the model may be used to predict an average melt pool temperature.

Results demonstrate that the model may be used to predict an average melt pool temperature. Bulk, calculated temperature distribution needs to be further investigated to find a more suitable heat transfer coefficient surrounding the part.IntroductionLaser Engineered Net Shaping (LENSTM) is a blown-powder, laser-based Directed Energy Deposition (DED) additive manufacturing (AM) process for metals fabrication originating from Sandia National Laboratories [1], [2]. In this process, a laser creates a molten pool as a CNC- guided substrate moves at a predefined speed and path, while metal powder is continuously injected into the pool to form a solid deposit/track. A part is thus fabricated track by track, and then layer by layer, in free space instead of within a powder bed. Due to the nature of its process, LENS offers the ability to control microstructure [3]–[5], fabricate parts of higher strength [6]– [9], and create functionally graded parts [10], [11].

Due to the nature of its process, LENS offers the ability to control microstructure [3]–[5], fabricate parts of higher strength [6]– [9], and create functionally graded parts [10], [11]. The challenge still exists, however, to adequately understand and model thermal aspects of the process in order to best select the appropriate process parameters and to more confidently qualify fabricated parts for use in the field. The process-property-performance relationships inherent to LENS and its produced materials require one to relate process parameters to part thermal response during manufacture; allowing one to predict final microstructure and residual stress.350To determine the thermal response of LENS materials during manufacture, one may resort to numerical simulation or analytical modeling to solve the heat equation representing heat transfer in successive layers.

In addition to thermal exchange with surroundings (including laser and chamber gas/walls) via radiation and convection, melt pool formation and solidification, as well as its fluid motion, can also be considered at varying levels of fidelity. Analytical solutions are desirable in that they offer an elegant representation of the thermal phenomena, e.g. cooling rates and temperature gradients, inherent to the process. Perhaps the most well-known of these is Rosenthal’s solution to the 3D heat equation for a moving point heat source [12]. Originally intended for welding, the solution has since been applied to DED due to process similarities and has since aided the understanding of melt pool geometry and solidified microstructure for various AM process parameters [3], [13], [14].

Neither latent heat of fusion effects nor temperature- dependent properties are taken into consideration through Rosenthal’s approach; and several researchers have noted the need for such information to be integrated for obtaining more accurate solutions [13], [15], [16].Applying welding models to the DED process has challenges. These models tend to not include mass addition to the melt pool or the laser attenuation created by the DED powder flow. Welding may also deviate from the DED process when deposited surface is considered. For instance, many welded surfaces are large enough that the deposition plane may be considered infinite. In DED, however, a track could be deposited on such a plane where edge effects may need to be considered. A more recent solution to the DED process was sought by Gockel et al. [17] to describe these edge effects.

A more recent solution to the DED process was sought by Gockel et al. [17] to describe these edge effects. They solved the heat equation for edge effects on a thin wall using the superposition of two point heat sources symmetric about the edge of the wall. Though thermal properties were temperature-independent, the solution to the melt pool geometry trended well with an FEA solution with temperature-dependent properties. Others have also formulated analytical models for various phenomena in DED such as clad geometry, beam attenuation, and powder heating that may be paired with other models for increased accuracy of results [18], [19].Process maps (i.e. surface plots) have been developed from Rosenthal’s solution for aiding process parameter selection [13], [14]. Process maps, in particular, show great potential in guiding the process parameter selection by describing melt pool and temperature response with a change in dimensionless operating parameters.

Process maps, in particular, show great potential in guiding the process parameter selection by describing melt pool and temperature response with a change in dimensionless operating parameters. The mapping scheme is realized by solving Rosenthal’s equation with changing variables and using the temperature solution to determine thermal gradients and cooling rates. These values are then used to study effects on melt pool geometry, residual stress, and solidification microstructure. However, it can be difficult to incorporate different part geometries into Rosenthal-based maps since such solutions are based on parts with semi-infinite domains and no edge effects. A ‘scaling factor’ between geometries is desired.This current study focuses on establishing and examining an alternate approach for predicting DED process and material property/size effects on subsequent thermal response. A purely-analytical thermal resistance network modeling approach is adopted herein.

A purely-analytical thermal resistance network modeling approach is adopted herein. Via Ohm’s analogy, the complex heat transfer occurring during DLD can be approximated as electrical current with one-dimensional flow along a nodal circuit containing various capacitors and resistors. Using such an approach allows one to isolate each mode and avenue for heat transfer and allows for a more intuitive representation of the thermophysics. The method is a ‘piecewise’, low-fidelity analytical approach enabling the determination of various temperature differences and heat fluxes351while introducing reduced calculation time.

The method is a ‘piecewise’, low-fidelity analytical approach enabling the determination of various temperature differences and heat fluxes351while introducing reduced calculation time. Process parameters may then be estimated according to geometry to ensure products of a particular quality (microstructure, residual stress, etc.) are produced.Thermal Resistance Network ConfigurationModel DevelopmentThe schematic of the laser-based/blown-powder DED, or LENS, process, is shown in Figure 1 where a laser is directed at the center of the top layer of a deposited thin wall. Powder is fed into the melt pool created by the laser to form a weld bead in this model. During the actual DED process, the laser moves relative to the substrate in order to deposit a track of molten material rather than a bead.

During the actual DED process, the laser moves relative to the substrate in order to deposit a track of molten material rather than a bead. To simplify the problem, a static laser spot is studied at the center of deposited thin walls, which may represent a part’s geometry at an arbitrary time during the build, to determine the effects of a varied geometry on average melt pool temperature and average temperature distribution within the part. Assumptions and idealizations employed for the model follow many of those discussed by Pinkerton & Li [20] and are omitted for brevity. Additional assumptions include:LaserMetal PowderWeld BeadThin Walled PartMelt PoolHeat Affected ZoneSubstrateFigure 1. Schematic of DED system in consideration.1. The laser is not moving along part, and no preheat exists on the part’s surface. 2. The initial condition consists of the melt pool.

Schematic of DED system in consideration.1. The laser is not moving along part, and no preheat exists on the part’s surface. 2. The initial condition consists of the melt pool. That is, some volume of material is assumed to have already experienced a phase change to a liquidus state. This is a valid assumption since the tested processing variables is known to create parts through experimentation.3. Rosenthal’s closed-form solution [12] to the heat equation of a moving point heat source on an infinitely large substrate may be used to find the depth of the melt pool.4. Only the melt pool surface area is radiating to the environment. 5. Surrounding environment is an enclosed volume at room temperature significantly larger than the part being fabricated.3526. The thin wall is sufficiently tall such that its temperature at the bottom is in thermal equilibrium with surroundings.7.

The thin wall is sufficiently tall such that its temperature at the bottom is in thermal equilibrium with surroundings.7. The empirical correlation for the convection heat transfer coefficient in vicinity of the melt pool as devised by Heigel et al. [21] is valid.8. Single phase heat transfer within the melt pool may be modeled using an effective thermal conductivity as described by Lindgren [22].9. The part behaves as a fin during laser irradiation; i.e., one dimensional, hyperbolic conduction.The heat transfer initiating from the surface of the melt pool, originating from the incident laser, is determined by compensating for thermal exchange with the injected powder cloud and spectral limitations of interacting media. Pinkerton and Li’s model is employed here by making similar assumptions [20]. Laser power absorbed and reflected by particles within an inner region to the melt pool is given by Eqs. (1) - (2), respectively [20].

Laser power absorbed and reflected by particles within an inner region to the melt pool is given by Eqs. (1) - (2), respectively [20]. The blown-powder, laser-based DED of titanium alloy, Ti-6Al-4V, is chosen as the modeled scenario. The material properties and process parameter values used for calculations are given in Table 1 and Table 2, respectively, along with the symbols for each.𝑞ap1 =3ζp,Ti𝑚̇ 𝐷1𝑙b 𝜋𝑟powder𝑣powder𝜌s𝐷03 ∙ 𝑞input𝑞rp1 =3(1 − ζp,Ti)𝑚̇ 𝐷1𝑙b 𝜋𝑟powder𝑣powder𝜌s𝐷03 ∙ 𝑞inputThe solid absorption, ζp,Ti, is dependent upon laser type and has here been idealized at 1.064 μm (wavelength of common Nd:YAG lasers) and 473 K to take into account an expected heating of the powder. Power absorbed and reflected within an outer region to the melt pool is given by Eqs.

Power absorbed and reflected within an outer region to the melt pool is given by Eqs. (3) - (4), respectively.𝑞ap2 =3ζp,Ti𝑚̇ (𝐷0 − 𝐷1)𝑙b 𝜋𝑟powder𝑣powder𝜌s𝐷03 ∙ 𝑞input𝑞rp2 =3(1 − ζp,Ti)𝑚̇ (𝐷0−𝐷1)𝐿𝑏 3 𝜋𝑟powder𝑣powder𝜌s𝐷0𝑞inputThese output values are then combined to create a compensated power input (W) perPinkerton et al. [20], 𝑞I, by using Eq. (5).𝑞I = ζp,Ti(𝑞input − 𝑞ap1 − 𝑞ap2 − 𝑞rp1 − 𝑞rp2) + 𝑞ap1The compensated power is then used as the input energy in Rosenthal’s moving heat source equation, since it considers particle interactions before the laser powder is incident to the melt pool. Rosenthal is used to determine the depth of the melt pool, dmp, so that the melt pool may be modeled in the system as an oblate, ellipsoidal component whose diameter is determined by the laser beam, D1, and whose deposited height is equal to the depth.353(1)(2)(3)(4)(5)Table 1.

DED process parameters used for model and experimentation.Process Parameter ValueLaser Power (qinput) 300 WScan Speed (vlaser) 12.7 mm/sPowder Feed Rate (𝑚̇ ) 0.32 g/sLaser power is further compensated by subtracting energy absorbed by the melt pool’s latent heat of fusion; assuming that the melt pool is already liquidus as the initial condition. The overall process is modeled as a non-moving laser; however, energy lost to latent heating is a known function of relative laser scan speed. For instance, more energy would be lost to latent heating if the scan speed were increased, because less power would be supplied to an area due to small surface-laser interaction times. This lost power contributes toward keeping the melt pool in a liquid state. In order to devise a relationship between latent heat of fusion and lost power, the melt pool is represented as an ellipse perpendicular to the scan path moving at some velocity as shown in Figure 3.

In order to devise a relationship between latent heat of fusion and lost power, the melt pool is represented as an ellipse perpendicular to the scan path moving at some velocity as shown in Figure 3. The following equation for lost power to latent heating is then determined:𝑞melt = 𝐿 ∙ 𝜌s ∙ 2 ∙ 𝑑mp ∙ 𝐷1 ∙ 𝑣laser(6)Equation (6) describes the power required to maintain moving, molten metal mass described by the cross-sectional area of the melt pool shown in Figure 3. The value of Eq. (6) is then subtracted from the net heat flux inputted to yield a final, adjusted value for power input into the thermal circuit, qII.354Deposition planevlaserD1dmpFigure 3. Depiction of melt pool as it moves along a plane of deposition.Thermal Resistance ModelThe thermal resistance model considers the temperature response of a completed thin wall that has had time to cool to room temperature and is subjected to heating in the center of the top edge.

The latent/spectral/powder compensated, DED heat transfer through the melt pool and part, as well as with the environment, is modeled using a procedure similar to Figure 2 and a thermal resistance network shown in Figure 4. Ohm’s analogy likens the heat flow to electrical current through a circuit and temperature difference to voltage difference at nodes. Resistors represent an item that restricts the heat flow and can arise due to conduction, thermal spreading, contact area, etc. A material may also store heat energy in the same way a capacitor in an electrical circuit stores energy. The amount of heat flow through a resistor, qresist, and capacitor, qcap, is given by Eqs.

A material may also store heat energy in the same way a capacitor in an electrical circuit stores energy. The amount of heat flow through a resistor, qresist, and capacitor, qcap, is given by Eqs. (7) & (8), respectively.𝑞resist =∆𝑇 𝑅(7)𝑞cap = 𝐶𝑑𝑇 𝑑𝑡(8)where R represents any resistance, ΔT is the temperature difference between nodes on either side of a resistance, C is a material’s capacitance as355Input Material & System Properties and Process ParametersStartCalculate useful energy before surface interaction per PinkertonCalculate melt pool depth per RosenthalUse melt pool depth to determine power lost to latent heatCircuit component calculation from Eqs. (13), (14), & (16) Keep radiation resistance as function of TsSolve for Ts as a function of timeDisplay Ts as a function of timeInput desired time to view temperature distribution in secondsDisplay temperature distribution from top of part to bottomFigure 2.

Process chart to determine the temperature distribution within a part.determined by Eq. (14), and 𝑑𝑇 𝑑𝑡⁄ is a temperature difference overtime as experienced by the capacitor. A positive 𝑑𝑇 𝑑𝑡⁄ indicates that the capacitor is being charged, whereas a negative value indicates that the capacitor is being discharged.Figure 4. Thermal resistance network representation of heat transfer during the DED of a rectangular wallThe empirical, forced convection model developed by Heigel et al. [21] is employed to calculate a heat transfer coefficient across a thin wall, hwall. The correlation is given below:ℎwall = (−2.717𝑧 + 37.174) ∙ 𝑒−(0.107𝑟)2.7+ ℎ0wallwhere z is the distance from the powder supply nozzles to the top of the wall in mm, r is the distance from the jet centerline to a point of interest in mm, and h0wall is the free-convection coefficient near 25 W/m2∙K [21]. Free convection becomes dominant around r = 15 mm.

Free convection becomes dominant around r = 15 mm. In the overview schematic of Figure 4, this resistance is only connected to the surface temperature of the melt pool; hence, Eq. (9) is only integrated over r as the diameter of the melt pool. The value for z is held constant at lb as defined in Table 1.The temperature-dependent, radiation heat transfer coefficient, hσ(Ts), is described in Eq.(10):ℎσ(𝑇s) = 𝜀lσ(𝑇s + 𝑇∞)(𝑇s2 + 𝑇∞2)where σ is the Stefan-Boltzmann constant, 5.67∙10-8 W/m2∙K4. The heat transfer coefficients are then converted into thermal resistance via Ohm’s analogy and the following equations result:𝑅h = 1 ⁄ ℎwall𝐴mp𝑅σ = 1 ⁄ ℎrad(𝑇s)𝐴mp356(9)(10)(11)(12)where Amp is the surface are of the oblate, ellipsoidal melt pool with two diameters equal to D1 and the other equal to twice the dmp.

The resistances are then combined in parallel to yield Rh+σ by:−1 = 𝑅h 𝑅h+σ−1 + 𝑅σ−1The capacitance of the melt pool, Cmp, is dependent upon amount and properties of themelted material and is defined in Eq. (14):𝐶mp = 𝑐p,l𝜌l𝑉mpwhere Vmp is the volume of the melt pool. The melt pool is assigned an effective thermal conductivity, kmp; therefore, the resistance through the melt pool, Rmp, is calculated using a standard conductive resistance model:𝑅mp =𝑑mp⁄𝑘mp𝐴mpIn the thermal circuit shown Figure 4, the top node receives the compensated heat transfer input from the laser and represents the melt pool surface temperature. From this node, heat (analogous to current in an electrical circuit) is allowed to flow to the convective/radiative resistance (Rh + Rσ), melt pool capacitance (Cmp), and melt pool (Rmp). The addition of the three heat flows should add to give qII.

The addition of the three heat flows should add to give qII. Convective, radiative, and capacitive resistances were all modeled as to terminate at ambient temperature. For the capacitor, the environment determines the amount of energy storage available. The same amount of power going into the combined melt pool and spreading resistances will also travel through the bulk part. However, average temperature (analogous to voltage in an electrical circuit) of the melt pool surface, Ts, and melt pool bottom, Tmelt, will be dissimilar. The former value will be unknown while the latter is the material’s liquidus temperature. By knowing the input power, ambient temperature, and temperature at the bottom of the melt pool, the average surface temperature of the melt pool may be found.

By knowing the input power, ambient temperature, and temperature at the bottom of the melt pool, the average surface temperature of the melt pool may be found. Only an average temperature is found due to the one-dimensional nature of the system.Finally, the intra-part heat transfer after the melt pool was assumed to be representative of heat transfer from an extended surface. The hyperbolic equation used to describe the temperature distribution within a rectangular fin of prescribed end temperature is used. The equation has been excluded for brevity, but is well known and easily accessed [31]. Since the current setup is essentially a “reversed fin”, the “base temperature” (hot side) of the fin is taken as the average temperature of the upper surface. This distribution is solved while accounting for the convection around the part as well as the conduction through it.

This distribution is solved while accounting for the convection around the part as well as the conduction through it. Rosenthal’s solution has been utilized to determine the average temperature of the part’s upper surface between the edges of the part and melt pool. The value is used as the upper surface temperature in the fin temperature distribution. Average melt pool temperature is calculated directly from the resistance network and is included in the overall surface average.357(13)(14)(15)ExperimentsThe two thin walls of different lengths in Figure 5 were printed in order to validate the current model. Because the walls were thin, no value was prescribed for the wall thickness which was measured to be around 2 mm after fabrication. This thickness is largely a result of the selected process parameters used in Table 2 which were determined before hand through experimentation to yield fully dense thin walls.

This thickness is largely a result of the selected process parameters used in Table 2 which were determined before hand through experimentation to yield fully dense thin walls. Height was kept constant at 21.09 mm and length, a, was changed from 25.4 mm to 50.8 mm. An OPTOMEC LENSTM 750 system with a Nd:YAG laser was used, in conjunction with spherical Ti-6Al-4V particles. The molten metal was deposited track upon track to build up the thin walls atop a substrate of the same material by moving a build platform underneath the static laser.21.09 mm21.09 mm~2 mm~2 mma = 25.4 mma = 50.8 mmFigure 5. Dimensions of thin walls used in experiments to validate model.An infrared (IR) camera (calibrated to black-body temperature) mounted to the moving build platform gathered images of the bulk thermal histories of each wall. The IR camera possessed a spectral response range of 8–14 μm and was set to capture images at a rate of 14 Hz.

The IR camera possessed a spectral response range of 8–14 μm and was set to capture images at a rate of 14 Hz. Each pixel was found to correspond roughly to about 0.37 mm on the build. Because IR cameras are difficult to calibrate to true temperatures due to a materials emissivity’s dependence on view angle and spectrum, normalized IR values were used in model comparisons. To do this, each row of pixels extending from edge to edge along the length, a, of the wall was averaged together from the top of the wall to the bottom. An average temperature distribution along the height of the walls was thus obtained. The distribution was then normalized by dividing all of the average values by the maximum average value.

An average temperature distribution along the height of the walls was thus obtained. The distribution was then normalized by dividing all of the average values by the maximum average value. Each image at the temporal mid-point and end of the build was processed individually in the same fashion with the laser at the center of the upper surface.A dual-wavelength (DW) pyrometer was set up to view the melt pool at all times and captured images at a rate between 4 – 7 Hz. For each image, all temperature values exceeding Tmelt were averaged together to determine the experimental, average melt pool temperature. Melt pool images and the immediate heat affected zone (HAZ) are shown in Figure 6 (a) and (b) during the final layer of the small and long length builds, respectively. Any values greater than Tmelt belong to the melt pool while the rest are part of the HAZ (values < 1300 °C are considered noise358and have been excluded).

Any values greater than Tmelt belong to the melt pool while the rest are part of the HAZ (values < 1300 °C are considered noise358and have been excluded). The melt pools are shown to be of relatively the same size and shape, though (b) shows slightly higher average temperature values.(a)(b)Figure 6. Melt pool images taken as the laser deposited in the center of the part during the final layer of the a = 25.4 mm (a) and a = 50.8 mm (b) builds. Temperature scale in °C.Results & DiscussionThe calculated, average melt pool temperature using the thermal resistance network and experimentally-measured, average melt pool temperature from the two thin wall builds are shown together in Figure 7. As calculated, the average melt pool temperature was found to be 2051 K for both sized walls. In fact, the height of the wall was not found to affect this value within the ranges tested.

As calculated, the average melt pool temperature was found to be 2051 K for both sized walls. In fact, the height of the wall was not found to affect this value within the ranges tested. The calculated value translates to an error of 2% and 3% for the 25.4 mm and 50.8 mm tall walls respectively. Pyrometry data agree with the calculations to the extent that the highest sustained temperature of the melt pool is the same. One should recall that the calculations from the thermal network occurred with a still laser whose temperature was found to reach steady state around 2 s after being turned on. The experiment itself consisted of consecutive layers building upon each other, and the data it produced shows average melt pool temperature at every point during the build.

The experiment itself consisted of consecutive layers building upon each other, and the data it produced shows average melt pool temperature at every point during the build. Figure 7 shows that the wall of smaller a value starts at a certain temperature, steadily rises, and then achieves steady-state temperature similar in magnitude to which it began.Bulk temperature data was gathered by an in-chamber, infrared (IR) camera mounted on the moving stage to which the substrate was secured. The average temperature distribution was extracted from the top to the bottom of each thin wall geometry at the temporal mid-point of the build as well as the completion. Part temperature distribution was modeled as a fin with prescribed values for end temperatures as mentioned in the previous section. The results of this model represent the average temperature distribution along the height of the part.

The results of this model represent the average temperature distribution along the height of the part. To normalize these values, the points on the temperature distribution were divided by the maximum, calculated value.359Comparative results between normalized experimental and model data are shown in Figure 8 (a) and (b).2250) K ( e r u t a r e p m e T20001750Average Calculated Melt Pool TemperaturePyrometry Data for a = 25.4 mmPyrometry Data for a = 50.8 mm1500050100150200250Time (s)300350400Figure 7. Comparison of average melt pool temperature between the a = 25.4mm and a = 50.8mm thin walls against build time. Each pyrometry data point shows the average temperature of the melt pool at a given time. The calculated, average temperature was found via the resistance network.Figure 8 (a) shows the normalized, average temperature distribution along the thin wall of smaller length.

The calculated, average temperature was found via the resistance network.Figure 8 (a) shows the normalized, average temperature distribution along the thin wall of smaller length. As seen in the figure, values align at the top and bottom of the wall at both wall heights. This is, of course, expected since the top of the wall is being heated while the bottom is near ambient conditions. However, the modeled distribution predicts a linear decay as opposed to the exponential decay shown by the data. This suggests that the free convection, heat transfer coefficient used in the fin model is too low. Upon increasing this coefficient, the model begins to approach the experimental data. When the length of the part is increased, the model seems to decrease in accuracy. While an increased heat transfer coefficient will still cause the model to approach the data, the normalized endpoints remain incongruent.

While an increased heat transfer coefficient will still cause the model to approach the data, the normalized endpoints remain incongruent. This phenomenon may be related more to the experimental conditions than the modeled conditions. Depending on the relative location of the part to the IR camera, results could vary slightly due to the angular variance of a material’s emissivity.360450(a)1.2) K / K ( e r u t a r e p m e T10.80.60.421.09 mm13.69 mma = 25.4 mmTest: 13.69 mm Tall0.2Test: 21.09 mm Tall0Calculated: 13.69 mm Tall0510152025Calculated: 21.09 mm TallDistance from top to bottom of wall (mm)(b)1.2121.09 mm10.73 mm) K / K ( e r u t a r e p m e T0.80.60.4a = 50.8 mmTest: 10.73mm TallTest: 21.09 mm Tall0.2Calculated: 10.73 mm Tall0Calculated: 21.09 mm Tall0510152025Distance from top to bottom of wall (mm)Figure 8.

Test and calculation comparisons of a=25.4 mm (a) and a = 50.8 mm (b) thin wall temperature profile.Finally, the uncalibrated, bulk, average temperature distributions from the experimental IR data are compared in Figure 9. Higher surface temperature is seen in the smaller length thin walls than the larger – a trend that is mimicked by the fin-modeled distribution. This occurrence may not only be explained by the smaller dwell time between layers for the smaller build, but it may also be attributed to the edge-effects in the build. Since the part is smaller, more heat tends to remain near the edges: increasing average temperature. Furthermore, the longer build shows a lower relative temperature simply because there is a greater length of the build. Thus, the expanse of the lower temperatures is greater from edge to edge. Because of these observations, average temperature distribution within the part is shown to be dependent upon size.

Thus, the expanse of the lower temperatures is greater from edge to edge. Because of these observations, average temperature distribution within the part is shown to be dependent upon size. This conclusion has little merit in-and-of itself, but it does suggest that thermal gradients within the part will be different as size changes. Grain growth would then be indirectly affected through the average temperature distribution.361700600) s s e l t i n U500400a = 25.4 mm, 13.69 mm tall(la n g i S R300200a = 25.4 mm, 21.09 mm talla = 50.8 mm, 10.73 mm tallI100a = 50.8 mm, 21.09 mm tall0051015202530Distance from top to bottom of wall (mm)Figure 9. Comparison between the average temperature distribution of the thin walls of varying lengths shown near halfway (solid lines) and at the completion (dashed lines) of the builds.

Comparison between the average temperature distribution of the thin walls of varying lengths shown near halfway (solid lines) and at the completion (dashed lines) of the builds. The IR signal has not been calibrated and is therefore considered unitless.ConclusionsA thermal resistance network to determine the average surface temperature of the melt pool has been devised and validated against experimental data. The calculated surface temperature is in good agreement with the data, and does not appear to vary with increased wall length. However, further experiments are hypothesized to show a decreased average melt pool temperature with increased thickness. The model may aid in selecting preliminary parameters (particularly laser power and powder feed rate) to aim for a certain average melt pool temperature.

The model may aid in selecting preliminary parameters (particularly laser power and powder feed rate) to aim for a certain average melt pool temperature. This can be useful when minimizing energy input is desired.Average temperature along the height of the thin wall has also been idealized as a fin temperature distribution with Dirichlet boundary conditions. When compared to normalized, experimental data the idealized results seem to suggest that a higher heat transfer coefficient in the model would be required to approach the real-life trend. Average temperature distribution within the wall was shown to decrease with increased thin wall length by both model and data values. Though this does not directly correspond to a change in microstructure, it may suggest a change in thermal gradients that could affect grain growth direction.Model results may be improved by incorporating thermal resistances into the model.

Better understanding on the connection between thermal spreading and the melt pool will be explored in order to increase accuracy of the melt pool temperature prediction. Future research will also aim to solve the heat equation for thermal spreading with DED boundary conditions as well as a moving heat source. Such an equation will allow researchers to further understand the part’s size effect on melt pool and bulk temperature distributions.The thermal network presented is a step towards creating a tool that may be used by researchers as a preliminary to parameter selection when part size is changed. It is not meant to replace detailed FEA models, but is intended to service those who may not have access to large, data-handling hardware or FEA software. This model is also not meant to replace process maps that present relationships between process parameters (laser power, scan speed, etc.) and part362properties (solidification microstructure, residual stresses, etc.).

However, it aims to supplement these maps by providing thermal information on detailed geometries.AcknowledgementsResearch was partially sponsored by the Army Research Laboratory and was accomplished under Cooperative Agreement Number W911NF-15-2-0025. The views and conclusions contained in this document are those of the authors and should not be interpreted as representing the official policies, either expressed or implied, of the Army Research Laboratory or the U.S. Government. The U.S. Government is authorized to reproduce and distribute reprints for Government purposes notwithstanding any copyright notation herein. This manuscript was prepared while Nima Shamsaei, Scott Thompson and Steve R. Daniewicz were faculty at Mississippi State University.References[1] M. L. Griffith, D. M. Keicher, C. L. Atwood, J. A. Romero, J. E. Smugeresky, L. D. Harwell, and D. L.

viewed PaperEVALUATING CONCEPTS FOR THE INTEGRATION OF MILLED COMPONENTS INTO THE ADDITIVE MANUFACTURING PROCESSJ. Reichwein*, J. Geis*, and E. Kirchner*Department of Product Development and Machine Elements, Technical University Darmstadt, GermanyAbstractLaser Powder Bed Fusion (L-PBF) has specific advantages over conventional manufacturing processes. These include high freedom in the design of components and cost- efficient production of small quantities. However, the surface quality of components is low compared to milling and the production of large components is often associated with high costs. These challenges are addressed by integrating milled components into the L-PBF process. Therefore, various concepts are presented for positioning, aligning, and fastening machined components in the build space of the L-PBF system with the goal to provide a reliable way to start the L-PBF process on top of these components.

Thus, allowing the potential of additive and subtractive manufacturing to be exploited without requiring an additional joining operation. Finally, these concepts are applied to a steering shaft bracket and the costs for manufacturing are evaluated. A 25% reduction in manufacturing costs was achieved compared to the purely additively manufactured component.IntroductionAdditive manufacturing offers great potential for geometric design through the layer-by- layer fabrication of components, thereby improving products in various applications. For example, the design freedom of additive manufacturing can be used to reduce mass and provide a more efficient use of a product for example by reducing fuel consumption [1]. An even greater reduction in mass is possible when multiple components are integrated into a single one.

An even greater reduction in mass is possible when multiple components are integrated into a single one. However, the complexity of the integrated component increases in terms of the number of geometric features and thus the manufacturing costs, due to a higher need for support structures, post-processing and longer production time [2]. In the context of Design for Additive Manufacturing (DfAM), there are already approaches for merging multiple components into one. The separation of components for additive manufacturing for example, to reduce the quantity of support structures or to manufacture components that exceed the building space, is discussed less frequently in the literature on DfAM [3]. The main approach of these contributions is to perform a methodical separation of components that exceed the build space of the additive manufacturing system and can subsequently be manufactured separately.

However, component separation also allows the individual elements to be manufactured in different manufacturing processes, which will be referred to as hybrid manufacturing in the following discussion. In this way, the potential of both manufacturing processes can be profited from. A systematic procedure for the selection of components for hybrid manufacturing is presented in [4].The advantages of additive manufacturing are the high geometric freedom and the ability to produce complex geometries while producing less waste. In addition, the dependency of the1286production costs regarding the number of components is significantly lower than with machining or casting processes [5]. This enables the individualization of components for each single customer. On the other hand, the surface quality of components is low compared to machining processes and leads to long cycle times and, especially in the field of metal additive manufacturing, high costs [1].

On the other hand, the surface quality of components is low compared to machining processes and leads to long cycle times and, especially in the field of metal additive manufacturing, high costs [1]. Machining processes such as milling, for example, have significantly different properties. Here, a precise, repeatable process leads to good surface quality. Productivity is also higher than in additive manufacturing, for example, due to the lower cycle times. However, due to the subtractive nature of these manufacturing processes, a large amount of waste is produced depending on the geometry. The costs are strongly dependent on the number of pieces to be produced and individualization is only possible with higher effort.The combination of additive and subtractive manufacturing as shown in Figure 1 makes it possible to reduce costs and waste while improving surface quality in some areas.

The prerequisite for this is the adaptation of the design of the individual parts and the selection of the manufacturing process that is more suitable for the part in question. Design guidelines that support the designer in the selection of parts for hybrid manufacturing and the design of these parts have already been addressed in the literature [4].Additive ManufacturingLessmaterial wasteGeometricfreedomIndividualizationLong cycle-timePoor surfacefinishMillingRepeatablePreciseGoodsurfacefinishHigh productivityWastefulHybridFigure 1: Properties of additive manufacturing and milling, based on [6]A hybrid process, which combines the advantages of both processes and simultaneously cancels out the disadvantages, therefore seems beneficial. In addition, different materials comprising various properties can be bonded together.

In addition, different materials comprising various properties can be bonded together. For example, combinations of aluminum and titanium [7], aluminum and steel [8], and steel and copper [9] have already been investigated and have in some cases produced good results. In this way, potentials can be exploited that cannot be achieved either with additive manufacturing or with subtractive manufacturing alone. While a combination of additive manufacturing with other manufacturing processes is also conceivable, this contribution is limited to the consideration of milling and additive manufacturing using L-PBF. There are a few examples in literature in which milled components have been inserted into the build space [10].

There are a few examples in literature in which milled components have been inserted into the build space [10]. However, concepts for integration have not yet been developed systematically and a methodical comparison of the different solutions is missing.Concepts for the Integration of Milled ComponentsAs part of the development of new concepts, the function of inserting components into the build space of the additive manufacturing system is first divided into sub-functions. These sub- functions are placement, orientation and fastening of the components. According to Figure 2, the placement of the components corresponds to the x and y coordinates and the rotation around the z-axis. The placement is necessary so that the additively manufactured part of the component is1287built up at the correct position on the top of the milled component and no offset occurs between the two parts. The orientation describes the rotation of the part around the x- and y-axes.

The orientation describes the rotation of the part around the x- and y-axes. It is necessary to ensure parallelism between the recoater of the manufacturing system and the inserted component. Thus, at the start of the additive manufacturing process, a uniform first powder layer is achieved on top of the milled part. This is an important parameter for a solid fusion of the additively manufactured part to the milled component and thus a mechanically durable connection. Finally, it is often necessary to fasten the inserted components, since thermal expansion during exposure or contact with the coater can lead to displacement of the component during the additive manufacturing process.OrientationPlacementzyxFigure 2: Orientation and placement of the components on the build plate.The three sub-functions can be solved almost independently of each other. Different concepts for each of these sub-functions are therefore presented below.

Different concepts for each of these sub-functions are therefore presented below. The solutions for the sub- functions are then combined to several overall concepts. Finally, the various concepts are evaluated in terms of the integration effort required for one-off production and repeated production of the same component, and in terms of their suitability for the production of small and large components.PlacementThe placement of the milled components is important to avoid misalignment of the two differently manufactured parts. There are two basic options for placement: Machining the build platform so that the milled components can be inserted, as shown in Figure 3 a). The second option is the additive manufacturing of fixtures for positioning the components directly on the build platform (see Figure 3 b)). The machining of the build platform is initially more complex, but leads to a better fit and can be reused for the same component.

The machining of the build platform is initially more complex, but leads to a better fit and can be reused for the same component. For high volumes, the machining of the build platform is therefore preferable. If only small quantities are required, additively manufactured mounts are more suitable, as the build platform can also be reused for other components afterwards.a)b)Figure 3: Concepts for placing milled parts into the additive manufacturing process with a) machined build platform insertion area and b) additive manufactured fixture to be placed on top of build platform.1288OrientationIn addition to the placement of the milled components in the build space of the additive manufacturing system, an orientation of the components must also take place. The goal of this orientation is to ensure the surface on which the additive manufacturing process is started and the coater of the manufacturing system are parallel.

The goal of this orientation is to ensure the surface on which the additive manufacturing process is started and the coater of the manufacturing system are parallel. Lack of parallelism between the component and the coater can lead to an uneven first powder layer and thus possibly to poor bonding, and to an angular misalignment within the component. Therefore, the alignment of the inserted milled parts is mandatory to achieve good quality.Some machines are adjustable to perform this orientation for the build platform and thus also for the milled parts mounted on it. If this option is not available, care must be taken to ensure parallelism between the coater and the component when inserting the milled parts. Overall, the orientation of the components is still a challenge because the additive manufacturing systems currently available on the market are not designed for inserting milled components.

Overall, the orientation of the components is still a challenge because the additive manufacturing systems currently available on the market are not designed for inserting milled components. This means that a complex manual alignment of the individual components is required, especially in the case of multiple components.FasteningOnce the part has been placed and oriented in the build space, the part must be fastened in the current position. Form-fit, force-fit or material-fit principles can be used for fastening. Material bonds include adhesive or welded joints. The latter are excluded in advance due to the destructive detachability of the connection. As shown in Figure 4 a) adhesive bonds are possible, but the temperature resistance must be checked when selecting the adhesive. Depending on the adhesive, the advantage of adhesive bonding can be improved heat conduction from the milled component to the building platform.

Depending on the adhesive, the advantage of adhesive bonding can be improved heat conduction from the milled component to the building platform. Disadvantages can arise from mixing the adhesive and unmelted powder. For example, adhesive residues must be conscientiously removed after the additive manufacturing process. In addition, the component must be freed from adhesive residues and cleaned.BGravitationalforceMagneticforcegSpring pinSheet metaldownholdera)b)c)d)e)f)AdhesivebondScrewconnectionFigure 4: Concepts for fastening milled parts in the additive manufacturing process.1289Force-fit connections include frictional connections and magnetic connections (see Figure 4 b)). An example of a frictional connection by means of a longitudinally slotted spring pin is shown in Figure 4 c). For this purpose, holes must be provided in the holder and in the milled part itself into which the spring pin is inserted.

For this purpose, holes must be provided in the holder and in the milled part itself into which the spring pin is inserted. In the relaxed state, the spring pin has a larger diameter than the diameter of the bore. In the assembled state, the expansion of the pin causes a frictional force on the outer diameter, which connects the components to each other. The use of magnets requires a magnetic material, but also results in a magnetic attraction of the powder. This solution is therefore particularly suitable when a magnetic milled part is processed with a non-magnetic powder.An intermediate solution between force and form-fit connections is the screw connection as shown in Figure 4 d). This also requires holes in the milled part and a threaded hole in the fixture. The advantage of the screw connection is that it can be easily disassembled after the manufacturing process, and the force with which the component is fastened can be adjusted.

The advantage of the screw connection is that it can be easily disassembled after the manufacturing process, and the force with which the component is fastened can be adjusted. The disadvantage is the necessity to modify the milled part. Particularly for small components where low thermal distortion is to be expected, fastening can be dispensed completely if necessary. The insertion of the milled semi-finished products into the holders and the weight force are then sufficient to prevent displacement during the manufacturing process as shown in Figure 4 e). In particular, if many components are to be inserted, a metal sheet can be used as a hold-down device, as shown in Figure 4 f). The advantage is that many components can be fixed at once. In addition, the individual components do not have to be adapted and, for example, there is no need to drill a hole in the components.

The advantage is that many components can be fixed at once. In addition, the individual components do not have to be adapted and, for example, there is no need to drill a hole in the components. Furthermore, the powder can be spread only on top of the hold-down sheet, so that less powder is required to fill the build space to the first layer of the L-PBF process. The disadvantage is that it is necessary to manufacture a suitable metal sheet to serve as a hold-down device. This solution is therefore particularly suitable for the production of a large number of components.The various concepts offered differ in terms of the effort required to integrate milled components and the strength of the fastening of the components. These two criteria are therefore used to evaluate the concepts. In addition, a distinction must be made between the one-time effort, such as manipulation of the building platform, and a recurring effort, such as the application of an adhesive bond.

In addition, a distinction must be made between the one-time effort, such as manipulation of the building platform, and a recurring effort, such as the application of an adhesive bond. Finally, this leads to an evaluation of the quantities for which the respective concept is suitable. For large quantities, a high one-time effort can be accepted, whereas for smaller quantities, a lower recurring effort should be preferred.The second criterion is the strength of the attachment of the components to the building platform. For example, a bolted connection can transmit more force than an adhesive connection. The strength of the connection is particularly relevant for components that are expected to have a large thermal distortion. The warpage of the components can cause the components to move out of the mount. In addition, the warpage can cause the coater to collide with the component when applying new material, thus exerting a force on the component.

In addition, the warpage can cause the coater to collide with the component when applying new material, thus exerting a force on the component. The fixture must then be strong enough to hold the component in place. Thermal distortion occurs in particular with large exposed areas, i.e. when the components comprise a large volume. Accordingly, the different concepts can be evaluated on the basis of the size of the components.1290Part countComponentdimensionssmallbigsmallbigFigure 5: Comparison of the different concepts for integration of milled parts in the additive manufacturing process according to parts size and part count of a production run.EvaluationThe concepts developed above are evaluated here using an exemplary component from a racing vehicle. The component selected for additive manufacturing is the vehicle's steering shaft bracket. This is bolted to the carbon-fiber-reinforced polymer body and includes two roller bearings that hold the steering shaft.

This is bolted to the carbon-fiber-reinforced polymer body and includes two roller bearings that hold the steering shaft. Accordingly, some requirements are placed on the steering shaft bracket. For example, the component should be as light as possible, whereby the angular deviation of the steering shaft in operation must not exceed 0.4°. To meet these requirements, a topology optimization of the component was carried out to achieve the required stiffness with minimum mass. The resulting complex structure is particularly well suited for additive manufacturing, especially since the steering shaft bracket is only produced in small quantities. The geometry of the component is not limited by restrictions of the manufacturing processes but can be designed to suit the lightweight design based on the load.In contrast, however, there are requirements for the two bearing seats of the component.

Here, good surface quality with low tolerance is necessary to ensure the assembly and reliable operation of the bearings. These requirements cannot be met by additive manufacturing using L- PBF. In the previous approach, this meant that the additively manufactured components had to be reworked using machining processes in order to be able to produce the bearing seats in the corresponding quality. Here, as shown in Figure 6 the component is separated into two parts and one is manufactured additively and the other by machining. The lower part, which contains the bearing seats, is produced by milling, while the topology-optimized structure is produced additively. The two components are joined during the additive manufacturing process, which starts on top of the machined component.1291Figure 6: Separation of the steering shaft bracket into a milled part (blue) and an additively manufactured part.The steering shaft bracket is separated using a plane.

This surface later forms the base surface for the start of the additive manufacturing process. The steering shaft bracket is only manufactured in small quantities. Manipulating the build platform is therefore time-consuming and means that the build platform cannot be reused for other components. The milled component is aligned by means of an additively manufactured structure that represents the negative of the geometry of the bearing seats. For fastening the milled part of the steering shaft bracket, a clamping connection was first provided by means of a spring pin. For this purpose, a hole is provided in the milled part and the additively manufactured fixture, the spring pin is then first fastened in the milled part and then in the additively manufactured fixture when the component is inserted. However, during the first test, the inserted part was moved by contact with the coater, so the manufacturing process had to be aborted.

However, during the first test, the inserted part was moved by contact with the coater, so the manufacturing process had to be aborted. The fastening force by the spring pin is accordingly too low, so the design of the fixture was changed to a screw connection instead.2. Inserting, orienting, andfixingthemilledpart3. Additive manufacturingoftheremainingpart1. Additive manufacturingofthefixturesFigure 7: Concept for integrating milled parts into the additive manufacturing process.Figure 7 shows the planned procedure for integrating a milled component into the additive manufacturing process. First, the fixtures are additively manufactured directly on the build platform, which is then used to place, orient, and fasten the milled part, which is inserted in the1292second step.

First, the fixtures are additively manufactured directly on the build platform, which is then used to place, orient, and fasten the milled part, which is inserted in the1292second step. Finally, another additive manufacturing process is started on the flat upper side of the milled part, in which the remaining part of the component is then placed directly on the milled part.Figure 8 shows the steps of the realized manufacturing process. After the fixtures have been manufactured, the unfused powder must be at least partially removed from the build space of the additive manufacturing system (see Figure 8 a)). This then allows the insertion of the milled part. Figure 8 b) shows the fastening of the milled part by means of a screw connection in the build space of the system. Before the part is fixed, the surface must be aligned parallel to the recoater of the manufacturing system.

Before the part is fixed, the surface must be aligned parallel to the recoater of the manufacturing system. The success of the alignment can be seen from the uniformly thin layer of powder on the top of the milled part in Figure 8 c).a)b)c)Figure 8: Integration of a milled part into the additive manufacturing process.The final result of the hybrid manufacturing process is shown in Figure 9. The steering shaft bracket comprises the topology-optimized, additively manufactured structure and the milled bearing seats, which can be used without any further post-processing. Table 1 also shows the costs for hybrid production in comparison with purely additive manufacturing of the steering shaft bracket.

Table 1 also shows the costs for hybrid production in comparison with purely additive manufacturing of the steering shaft bracket. By inserting a milled component, the costs were reduced by 25 %.a)b)Figure 9: Final hybrid manufactured steering shaft bracket.1293Table 1: Comparison of the cost of additive manufacturing of the steering shaft bracket with hybrid manufacturing.Additive manufacturingHybrid manufacturingAdditive manufacturing process1,400 €909 €Milling process137 €Rework of the bearing seats46 €Sum1,446 €1,046 €Summary and OutlookIn summary, several concepts for the integration of machined components into the additive manufacturing process were developed and compared to each other with regard to the size of the components and the number of pieces to be manufactured. This enables to systematically select a suitable concept for the respective application.

This enables to systematically select a suitable concept for the respective application. Two of these concepts were also applied to an example component, the steering shaft bracket of a racing car. It was shown that a frictional connection using a spring pin could not absorb sufficiently high forces, which is why a screw connection was used in a second step. This enabled the successful hybrid production of the steering shaft bracket. A comparison of the costs also showed that hybrid production is 25% cheaper than pure additive manufacturing. This can be attributed in particular to the reduced expense for reworking the bearing seats.The orientation of the milled components is particularly challenging, as this currently requires manual intervention by the operator. Further concepts can be developed here in the future, but these will require to adapt the additive manufacturing system.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEVALUATING THE COMPRESSIVE STRENGTH OF AM RISERS FOR GREEN SAND METALCASTINGC. R. Hasbrouck1∗ and Samantha A. Melnik1 1Department of Industrial and Manufacturing Engineering The Pennsylvania State University, University Park, PA 16802AbstractWhile many metalcasting foundries have experimented with using additive manufacturing (AM) for patternmaking, the compressive strength of the tapered AM risers for green sand metalcasting has not yet been explored. This study investigates the effects of infill pattern type, infill density, and shell thickness on the compressive strength of a standard 3-inch diameter by 5-inch tall by 3-degree taper cylindrical riser manufactured with PLA using a material extrusion process.

The findings for these AM risers include plots and mathematical models of compressive strengths at three different scales of the standard geometry (full, three-quarters, and half), predicted build times and masses using other common infill patterns, potential failure mechanisms during use of AM and conventionally manufactured riser patterns, and considerations on design for both additive manufacturing and green sand metalcasting. It is concluded that AM risers can be incorporated into and perform well as part of conventional green sand metalcasting patterns.Keywords: additive manufacturing, material extrusion, metalcasting, compressive strength, risers, patterns, hybrid manufacturing, desktop printersIntroductionMetalcasting has spanned centuries of practice and is the backbone of the world’s industrial materials industry, with 90% of consumer goods containing cast metal [1].

Through the centuries, metalcasting has seen improvements in efficiency, mechanical properties, cost, part complexity, etc., all the while retaining significant aspects of traditional processes [2]. It has been shown that the appropriate leveraging of additive manufacturing (AM) processes can lead to the improvement of the metalcasting industry, allowing for increased part complexity, reduced lead times, and better prototyping [3].It is important to have knowledge of traditional metalcasting processes to understand where AM can be beneficial. There are two overarching moldmaking categories of metalcasting: expendable mold and permanent mold processes [4]. Each moldmaking process category has advantages and disadvantages, with the major difference between the two types being the retention of the mold post-casting. In the case of expendable mold processes, the mold is single-use and is either destroyed from functionality or discarded completely post-casting [5].

In the case of expendable mold processes, the mold is single-use and is either destroyed from functionality or discarded completely post-casting [5]. Likewise, there are two overarching patternmaking categories of metalcasting: expendable pattern and permanent pattern processes [5-8]. Similarly to expendable molds, expendable patterns are considered as single-use while permanent patterns are reused for multiple∗Corresponding author: crh33@psu.edu1847castings. However, unlike expendable molds, expendable patterns such as foam and polymer- based materials are completely consumed during the casting process and do not leave any components behind that can be reused for a future expendable pattern [6].The sand casting process has been in operation since 645 B.C. and remains the cheapest and most widely used metalcasting process, continuing to be a widespread practice in a multitude of manufacturing sectors [2], [9].

and remains the cheapest and most widely used metalcasting process, continuing to be a widespread practice in a multitude of manufacturing sectors [2], [9]. Sand casting is an expendable mold metalcasting process that can utilize either permanent or expendable patterns [6], [7]. The expendable sand molds are typically categorized as green sand (silica or a specialty sand bonded with clay and water), or chemically bonded sand which uses a chemical binder [6]. If the mold is made from green or chemically bonded sand, it can be thermally or mechanically reclaimed for re-use as a new expendable mold or discarded entirely if it has lost functionality [6].In addition to traditionally manufactured sand molds, there are opportunities for additive manufacturing of sand molds via binder jetting processes as explored by Sivarupan et al. [10].

[10]. Companies such as ExOne and Voxeljet advertise such capabilities, offering printers with build sizes of 4 m x 2 m x 1 m, layer thicknesses as small as 0.15 mm, and print speeds exceeding 130 L/h [11-12]. Currently, these systems are costly, make use of hazardous binders, and require significant maintenance, thus the integration of such systems within the metalcasting community has been slow. However, Lynch et al. have explored the concept of shared industrial hubs for the use of binder jetting equipment, reducing the burden put upon a single metalcasting company, many of which are small businesses lacking the capital to leverage such technology otherwise [13].

Such industrial hubs for direct production of AM sand molds have not yet been implemented on a widespread basis, but the sand casting community can still currently make use of AM technologies for patternmaking.It is well known within the sand casting community that patternmaking is a considerable capital investment with substantial lead times for manufacture [14-16]. According to a metalcasting foundry local to the authors, the typical cost of a wooden pattern for a 26-inch by 26-inch by 8-inch flask is approximately $1600, and a similar pattern made from foundry board is about $5000; each has a typical lead time of about two weeks to produce [17]. If a pattern is insufficient in strength, dimensional accuracy, or results in unsound castings, it must be redesigned and remanufactured, further adding to lead times and capital costs.

If a pattern is insufficient in strength, dimensional accuracy, or results in unsound castings, it must be redesigned and remanufactured, further adding to lead times and capital costs. The ability to rapidly prototype or manufacture patterns using additive manufacturing reduces both lead time and capital costs by a considerable margin [15].As previously mentioned, the expendable mold sand casting process can use either expendable or permanent patterns. Permanent sand casting patterns are often machined from aluminum, wood, or foundry board while expendable patterns are traditionally made of injection molded foam or a polymer [7]. The use of desktop material extrusion processes has led to the in-between category of semi-permanent patterns; these patterns function as permanent patterns but due to their relatively low percentage of solid infill, they tend to either degrade or break more quickly than aluminum or wooden permanent patterns.

Such AM patterns are ideal for prototyping new design ideas or for low volume runs of castings [18].1848Prior work by Hasbrouck et al. has investigated the dimensional tolerancing capabilities of material extrusion-based patternmaking for green sand casting, including the use of polylactic acid (PLA) for non-critical pattern components and photopolymers for critical pattern components, but did not explore the strength of the patterns [19]. Anakhu et al. studied the effects of infill density, print layer height, and number of solid layers for optimum PLA plastic compressive strength [20], which is an important metric for determining the ability of the pattern to survive the compaction forces of sand required for moldmaking studied additively manufactured PLA’s in green sand casting processes. Pernet et al.

Pernet et al. compressive strength while varying the infill styles and percentages within the ASTM Standard D695-15 [21] test geometry, concluding that 2D infill patterns perform the most favorably when comparing load-to-weight ratios because of the alignment of the mechanical supports to the direction of the load. Harshit et al. performed a similar study of the compressive strength of additively manufactured PLA using ASTM Standard D695-15 [22], investigating the effects of layer height, infill density, and print speed on part strength and demonstrating optimal strength at a layer height of 0.2 mm, infill density of 80%, and print speed of 40 mm/min.

There are many other studies [23-33] which analyzed the effects on various mechanical properties of material extrusion-produced materials such as acrylonitrile butadiene styrene (ABS), polycaprolactone (PCL), polycarbonate (PC), PLA, PLA with an acrylic modifier, and ULTEM 9085 when modifying AM variables such as infill density, infill width, build orientation, laydown pattern, layer thickness, color, nozzle temperature, shell perimeters, and infill patterns.This case study focuses on the advancement of desktop material extrusion AM processes for nonferrous sand casting operations for the production of semi-permanent reusable patterns in the patternmaking process. There is a national interest in reshoring and revamping manufacturing capabilities in the United States by investing in development and integration of cutting edge technologies, as evidenced by state initiatives such as the Manufacturing PA Initiative in Pennsylvania [34].

In this vein, a nonferrous green sand metalcasting foundry, Erie Bronze and Aluminum Company (“Erie Bronze”), has experimented with AM for patternmaking in the past but gave up after the parts did not reduce lead time or meet strength requirements. The company expressed continued interest in the idea but needed assistance in determining printing parameters that resulted in a reduction in lead time while still meeting the minimum strength requirements for the forces imparted to the patterns by their molding practices. Erie Bronze uses both automated (Figure 1, (a)) and manual (Figure 1, (b)) green sand molding lines, each of which result in different force magnitudes and distributions across the patterns.Erie Bronze is a relatively small foundry that mainly functions as a ”job shop” with typical run quantities of as low as one and as high as 1000 parts [17], making it an ideal candidate for AM semi-permanent patterns.

The automatic molding line outputs a pressure reading of 700 psi [35], but it is unclear whether that is the hydraulic cylinder pressure, the pressure the flask sees overall, or the pressure the pattern parts see. Therefore, instead of trying to optimize the patterns for their ideal strength for this single application, the research presented in this study focuses on compressive strengths obtainable across a wide range of infill densities.1849Figure 1 : Molding lines at Erie Bronze and Aluminum Company: (a) automatic, and (b) manual. Published with permission.All green sand metalcasting patterns must contain a sacrificial rigging system to supply liquid metal to the part to be cast [5], [8]. Most metalcasting rigging systems contain at least one sacrificial volume of metal called a riser to feed the shrinkage that happens as the metal experiences a phase change during solidification from liquid to solid [36].

These risers are designed to meet solidification rules dependent on such variables as alloy shrinkage percentages, the volume of the casting to feed, feeding distance, thermal properties of the mold, and more [37] and are optimized to solidify after the rest of the casting system components. Therefore, the larger the volume-to-surface area ratio is according to Chvorinov’s rule [38], the better the riser due to the maximization of volume of metal coupled with the minimization of lowest cooling rates from conduction across the mold surfaces. Based on Chivornov’s rule, the ideal riser shape is a sphere. However, since spheres are difficult to mold for green sand patterns, tapered cylinders are most typically used instead. As these volumes of metal are sacrificial, their dimensional tolerances are less stringent than the functional surfaces of the actual cast part(s).

As these volumes of metal are sacrificial, their dimensional tolerances are less stringent than the functional surfaces of the actual cast part(s). Combining this knowledge with the relatively standardized riser geometry and frequency of its use across sand casting patterns led the authors to focus on risers for this case study. A nominal riser size of 3-inch diameter by 5-inch height by 3° taper as provided by Erie Bronze was used as a basis [17], [19]. An example of a matchplate featuring an AM semi-permanent pattern cope (a) and drag (b) used for green sand casting is shown in Figure 2, where the parts of interest are identified by blue ovals and the risers are indicated by orange arrows [19].1850Figure 2 : Matchplate (a) cope – top half, and (b) drag – bottom half, for green sand castings featuring additively manufactured patterns [19].The work for this study was completed with input from Erie Bronze to best solve their particular production issues.

The remainder of the paper documents the research questions explored, the experimental methodology used, and the results of the study.Research QuestionsFive research questions (RQ) and associated hypotheses (H) were developed to helpmetalcasting patternmakers design for additive manufacturing of risers:RQ1 : Does infill density (percentage) affect the print time, mass, compressive strength, dimensional stability, or surface roughness of additively manufactured risers of a standard length-to-diameter ratio?H1 : The infill density (percentage) will have a significant effect on print time, mass, andcompressive strength but not on dimensional stability or surface roughness.RQ2 : Does shell thickness affect the print time, mass, compressive strength, dimensional stability, or surface roughness of additively manufactured risers of a standard length-to-diameter ratio?H2 : The shell thickness will have a significant effect on compressive strength and surfaceroughness but not on print time, mass, or dimensional stability.RQ3 : Does infill type (pattern) affect the print time, mass, compressive strength, dimensional stability, or surface roughness of additively manufactured risers of a standard length-to-diameter ratio?H3 : The infill type (pattern) will have a significant effect on print time, mass, and1851compressive strength but not on dimensional stability or surface roughness.RQ4 : Does filament color affect the print time, mass, compressive strength, dimensional stability, or surface roughness of additively manufactured risers of a standard length-to-diameter ratio?H4 : Filament color will have a significant effect on compressive strength but not on printtime, mass, dimensional stability, or surface roughness.RQ5 : What are the print time, mass, compressive strength, and dimensional stability capabilities at various infill densities of additively manufactured risers of a standard length-to- diameter ratio for a constant infill type, shell thickness, layer height, and filament color?H5 : The print time, mass, and compressive strength will all increase with increased infill densities across all scales.

Dimensional stability will decrease as scale decreases but will not be affected significantly by infill density.Experimental MethodologyThis initial case study explores the use of material extrusion of polylactic acid (PLA) filament. Material extrusion was chosen due to its widespread availability, compact (desktop) nature, and relatively low capital, operating, and maintenance costs. PLA was chosen also for its widespread availability and low cost as well as its ease of printing and relatively good mechanical properties. Test specimens of a representative riser geometry were produced using two Prusa MK3S series printers. Infill pattern, density, and shell thickness were varied to test which combination provided optimal printing parameters.

Infill pattern, density, and shell thickness were varied to test which combination provided optimal printing parameters. Limiting cost and lead time are the two most important considerations for AM patterns for metalcasting, provided that the parts produced can withstand the forces of moldmaking and hold up dimensional tolerances to sand abrasion and moisture absorption over time. Thus, initial narrowing of parameters was done with that consideration in mind as to ensure significant lead time or cost was not incurred during the printing process, which would negate the benefit of its implementation. Upon narrowing down the printing parameters, parts were printed in three size scales to investigate the effects of part size on results.

Upon narrowing down the printing parameters, parts were printed in three size scales to investigate the effects of part size on results. Following printing, parts were compression tested to identify the yield strength of the various test specimens, to determine if AM risers were suitable for green sand operations.MaterialsDue to supply chain issues at the time of ordering filament, only a glitter filament was available to be procured in a quantity large enough for this study. A total of 15 kg of opal green Prusament PLA were ordered for this case study. It is important to note that the presence of glitter may affect both the mechanical and optical properties of the final parts depending on what is causing the glitter effect. For example, the presence of metallic powder [39] may cause the glitter effect, which can result in greater strength and further increases in brittleness [40] as well as increased nozzle wear [41].

For example, the presence of metallic powder [39] may cause the glitter effect, which can result in greater strength and further increases in brittleness [40] as well as increased nozzle wear [41]. While these parameters1852were not explicitly explored for this case study, it will be discussed how the presence of glitter negatively affected the ability to perform optical profilometry measurements on the printed parts.Fabrication and Physical MeasurementsTo begin, a nominal riser size of 3-inch diameter by 5-inch height by 3° taper as provided It was exported to a STL file using by Erie Bronze [17] was modeled in SolidWorks. American Standard Code for Information Interchange (ASCII) encoding, inch reference dimensions, and custom tolerance settings of 0.00032527-inch deviation tolerance and 0.5 degree angle tolerance. These settings provided the highest available quality STL file export from SolidWorks to minimize geometry losses to triangulation.

These settings provided the highest available quality STL file export from SolidWorks to minimize geometry losses to triangulation. Two different slicer softwares were used for this case study: Ultimaker Cura and PrusaSlicer. PrusaSlicer, unlike Ultimaker Cura, allowed for greater customization of part g-code when combined with the Prusa-branded printers; however, Ultimaker Cura was also employed, as it is one of the most commonly used open-source slicer softwares in the 3D printing community. Comparisons regarding print time and material requirements were made between the two softwares to ensure that the results were independent of the slicer software used. Two different pre-assembled Prusa material extrusion desktop printers were used: an Original Prusa i3 MK3S with orange trim, and an Original Prusa i3 MK3S+ with black trim.The printers were referred to by trim color instead of model number throughout data collection.

The height and bottom diameter of each printed part were measured using 8-inch Mitutoyo Digimatic digital calipers (Figure 3) with an accuracy of ± 0.0005 in (model number CD-8”ASX). For measurements of mass, a WeighMax digital postal scale with an accuracy of ± 3g (model number W-BC130) was used. To avoid user bias in the measurements, the same person performed all mass and dimensional measurements.Figure 3 : Measuring an AM riser with digital calipers: (a) height, and (b) diameter.1853DocumentationA print log was kept throughout the study that included the following information for every print: DOE number and quadrant or scale, shell thickness, infill pattern, infill density, print start time and date, print stop time and date, print pull time and date, initials of who pulled the print, the printer used, the filament roll ID number, the filament roll tolerance, and the filament roll manufacturing date.

An additional log for the measurements of each print was kept with the following information: DOE number and quadrant or scale, replication number, infill density, mass (g), print time (hr), height (in), diameter (in), initials of who measured the part, and the date the part was measured. Finally, on the bottom of the printed parts, identification of DOE number and quadrant or scale, shell thickness, infill pattern, infill density, print pull time and date, initials of who pulled the print, and the printer used were marked. Three replicates of each build parameter condition were printed on a single build plate.Surface RoughnessAny surface roughness measurements on AM parts were performed using a Zygo Zegageoptical profilometer shown in Figure 4.Figure 4 : Optical profilometry setup for AM risers.A 10X Mirau lens was calibrated with either 200 or 500 µm pitch square patterns on a standard Zygo calibration block.

A scan area/field of view of 830 µm by 830 µm was used1854with a scan depth of 1200 µm, resulting in a scan time of about 155 seconds for each part. Data was analyzed using Zygo Mx software employing conical form correction and ±3σ spike clip surface processing filters.Compression TestingCompression testing was carried out using a Tinius Olsen 120-kip universal testing machine (model number 120K Super L). Data was collected and initially plotted with the Tinius Olsen Test Navigator software. The machine was programmed to run a compression test from position data in four sequential steps:1. Move at a position rate of 0.5 in/min until a force value of 50 lbf is reached2. Zero channel force3. Zero channel position4. Move at a position rate of 0.1 in/min until a position of 1 inch is reachedStep 1 ensures the load cell is operating correctly and the part is touching both platens.

Zero channel position4. Move at a position rate of 0.1 in/min until a position of 1 inch is reachedStep 1 ensures the load cell is operating correctly and the part is touching both platens. Steps 2 and 3 clear the current force and position readings back to zero to perform the compression test. Step 4 is the compression test and was modified for each scale to stop at a position equal to 20% of the height of the sample or until a sample break was detected (75% or larger drop off in peak load), whichever came first. The data was automatically exported as a graphical image and as a comma-separated values (csv) file for both load (lbf ) and stress (psi) versus position (in). The software automatically reported yield strength based on the 0.2% offset rule.

The software automatically reported yield strength based on the 0.2% offset rule. Yield strength was chosen as the indicative strength over ultimate strength for two main reasons: (1) once the part has yielded to plastic deformation, the dimensions are no longer suitable for a metalcasting pattern, and (2) many parts never peaked or fractured at what could be considered an ultimate strength. Before each test, the sample name was inputted along with its height and the bottom (larger) diameter of the part.The compression testing setup is shown in Figure 5. Two test setups were run before the real testing to ensure the program ran as intended. First, two aluminum cans were crushed using the program to test the initial force readout and consequential zeroing of the channels as well as return of the crosshead upon reaching a position of one inch. Next, four dummy parts were printed using galaxy silver Prusa PLA, two at 15% cubic infill and two at 50% cubic infill.

Next, four dummy parts were printed using galaxy silver Prusa PLA, two at 15% cubic infill and two at 50% cubic infill. The dummy plastic parts were run to examine the sensitivity of the load cell, determine approximate low and high ends of strength of the parts, and determine possible failure mechanisms.Designs of ExperimentTo determine the best parameters to proceed with for the full experimental study, three designs of experiment (DOEs) were completed. All three DOEs used a two-level full factorial design with three replications and focused on the full-scale riser size with nominal dimensions. To reduce print time as much as possible, a large layer thickness of 0.3 mm was used and1855Figure 5 : Tinius Olsen 120-kip universal testing machine compression testing setup: (a) full view of machine, (b) loading a sample into the machine, and (c) data collection screen.held constant for all prints.

The first DOE aimed to answer RQ1 and RQ2 and investigated the significance of shell thickness and infill percentage on print time, mass, compressive strength, dimensional stability, and surface roughness. A gyroid infill was chosen and held constant, the high and low values for shell thickness were 1.4 and 0.5 mm, respectively and the high and low values for infill density were 50% and 15%, respectively. Similarly, the second DOE focused on RQ3 and investigated the significance of infill pattern on print time, mass, compressive strength, dimensional stability, and surface roughness.

Similarly, the second DOE focused on RQ3 and investigated the significance of infill pattern on print time, mass, compressive strength, dimensional stability, and surface roughness. A nominal shell thickness of 0.7 mm was chosen and held constant, the high and low values for infill pattern were gyroid and cubic, respectively (Figure 6) and the high and low values for infill density were 50% and 20%, respectively.Figure 6 : DOE infill patterns - (a) gyroid infill, and (b) cubic infill.1856Finally, the third DOE focused on RQ4 and investigated the significance of filament color on print time, mass, compressive strength, dimensional stability, and surface roughness. A nominal shell thickness of 0.7 mm and infill pattern of cubic were chosen and held constant, the high and low values for filament color were pineapple yellow and opal green, respectively and the high and low values for infill density were 50% and 20%, respectively.

Ultimaker Cura was used as the slicer for the DOEs because it provides additional manual settings for customization of shell thickness parameters over PrusaSlicer. All DOE data collected were analyzed using standard analyses of variance (ANOVAs).Full Experiments and Compression TestsUpon completion of the three preliminary DOEs, a full experimental design was developed to answer RQ5. The nominal riser dimensions were used and explored at three scales: full-scale, three-quarter scale, and half-scale. The layer thickness of 0.3 mm was continued to be held constant, as were the nominal shell thickness of 0.7 mm and filament color and brand of Prusament opal green PLA. A standard cubic infill pattern was chosen for the full experimental study due to its relatively high strength-to-weight ratio and significantly decreased print time over gyroid. Each scale of the nominal length-to-diameter ratio risers was completed with infill densities varying from 5-50% in increments of 5%.

Each scale of the nominal length-to-diameter ratio risers was completed with infill densities varying from 5-50% in increments of 5%. Each infill density had three replicates, all built on a single build plate. It is important to note that printing three replicates on a single build plate resulted in increased time between printing of layers than would normally be seen if printing one part at a time and could contribute to a slight decrease in mechanical properties. All full experimental data collected were analyzed using standard analyses of variance (ANOVAs) and linear regression modeling, where applicable.An attempt was made to replicate as closely as possible what metalcasting foundries would do if performing these experiments in-house. Therefore, time between print completion and compression testing was left relatively randomized and all replicates were built on a single build plate whenever possible.

Therefore, time between print completion and compression testing was left relatively randomized and all replicates were built on a single build plate whenever possible. Additionally, advanced design for additive manufacturing (DfAM) concepts such as topological optimization, customized infill lattices, and graded infill patterns were not explored.Wooden RisersAs discussed in the Experimental Methodology section, Erie Bronze typically uses wooden patterns for their automatic green sand molding machine. Since the 700 psi pressure reading provided by the company is not conclusive enough for designing an optimized plastic riser with infill pattern and density to reach the necessary strength, a generalized approach to determine achievable strengths with different DfAM parameters was employed instead.

Erie Bronze graciously supplied the authors with sets of pine wooden risers from patterns up to 20+ years old as well as new wooden risers ordered special from the pattern shop for this case study. The risers from the old patterns were at least partially coated for smoothness and contained additional pieces from gating system that were removed before testing. The wooden risers that were manufactured for this project were made to be the same scale as1857those plastic parts printed and tested. Examples at each scale of the old wooden risers, the newly manufactured wooden risers, and the plastic risers are shown in Figure 7.Figure 7 : Old wooden risers and new wooden risers next to their printed plastic counterparts at each of the three scales.Both the old and new risers consisted of between two to four pieces of wood assembled together, and the layer/grain orientation could be either vertical or horizontal. Examples of the different wood layer orientations are shown in Figure 8.

Analyses of variance (ANOVAs) at an alpha significance level of 0.05 were run for each parameter, considering the factors for the corresponding DOE.Four distinct datasets were produced during the data collection phase:DOE 1 – Infill Percent and Shell ThicknessThe significance of infill percent (low level: 15%, high level: 50%) and shell thickness (low level: 0.5 mm, high level: 1.4 mm) were studied in DOE 1. Results indicated neither shell thickness, infill percent, nor their interaction significantly affected riser height, diameter, or surface roughness. Surface roughness results were difficult to acquire optically due to glitter within the filament. Mechanical measurement techniques were attempted using a Mahr Pocket Surf IV surface roughness gauge; however, the roughness of the parts was beyond that of the equipment’s maximum measuring capabilities of 250 µin Ra.

Therefore, while such factors were not deemed significant for the surface roughness, additional work in this area is needed to verify the results.Infill percent, shell thickness, and their interaction significantly affected the parts’ compressive yield strength (lbf and psi), as well as the elastic modulus. Shell thickness was the only significant factor for break percentage. While shell thickness was found to1859significantly affect compressive yield strength (positively favoring 1.4 mm over 0.5 mm), visual inspection overrode the results. Both the high and low shell thickness values resulted in a vertical seam along the height of the cylindrical risers, pictured in Figure 9.Figure 9 : Vertical seams along PLA risers. 0.5 mm shell thickness (left); 1.4 mm shell thickness (right).While the risers are sacrificial parts, the seams alongside the risers would require additional post-processing to ensure a smooth finish or risk sand entrapment, adding extra time and cost.

Printed risers at a nominal shell thickness of 0.7 mm (standard for 0.3 mm layer height in Prusa Slicer) did not produce a vertical seam. Thus, the authors moved forward with the nominal shell thickness for future experiments.DOE 2 – Infill Percent and Infill TypeThe second DOE studied the effects of infill percent (low level: 20%, high level: 50%) and infill type (low level: cubic, high level: gyroid). Initially, DOE 2 was designed as a two-factor three-level full-factorial, with infill percent values of 20%, 50% and 80% and infill types gyroid, cubic, and honeycomb to replicate the case study published by Dev and Srivastava [43]. During the slicing process the experiment was redesigned, as infill percentages above 50% across all infill types began to eliminate the lead time advantages of AM over traditional patternmaking.

During the slicing process the experiment was redesigned, as infill percentages above 50% across all infill types began to eliminate the lead time advantages of AM over traditional patternmaking. The same reasoning led to the removal of the honeycomb infill pattern, which had a significantly longer print time across all infill percentages as compared to the cubic and gyroid infill patterns. Gyroid and cubic infill patterns were chosen as they are standard infill types found on most of the commercially available and open-source slicing software and have a favorable strength-to-weight ratio.Results from DOE 2 indicate infill type significantly affected riser diameter, whereas the interaction between infill type and infill percent significantly affected surface roughness. Though, similar to that of DOE 1, additional work is needed to verify the surface roughness1860results. Additionally, infill percent significantly affected the parts’ elastic moduli.

Though, similar to that of DOE 1, additional work is needed to verify the surface roughness1860results. Additionally, infill percent significantly affected the parts’ elastic moduli. percent, infill type, and their interaction significantly affected riser height, mass, and compressive yield strength (lbf and psi). At 20% infill, risers with gyroid infill were 0.57% stronger than parts with cubic infill. However, at 50% infill, parts with cubic infill were 2.6% stronger than those with gyroid infill. Given the substantially lower print times and mass requirements for the cubic infill, as well as the minimal difference in strength at the 50% infill mark, the authors chose to use cubic infill for the remaining experiments.DOE 3 – Infill Percent and Filament ColorThe third DOE focused on color type (low level: opal green, high level: pineapple yellow).

Due to supply chain issues and filament availability at the time of the study, the risers were initially printed in Prusament opal green PLA filament, which contains glitter. As discussed in the Experimental Methodology section, it is known that presence of glitter elements can affect mechanical properties of PLA. It has also been proven that the color of PLA filament can affect tensile strength of PLA parts [44]. To test the significance of PLA color on compressive strength, and in an attempt to acquire optical surface finish data, parts at 20% and 50% cubic infill were reprinted in Prusament pineapple yellow PLA to evaluate potential significance. Both infill percent, infill type, as well as their interaction were deemed significant factors affecting the parts’ compressive strength (lbf and psi), with yellow found to be favorable.

Both infill percent, infill type, as well as their interaction were deemed significant factors affecting the parts’ compressive strength (lbf and psi), with yellow found to be favorable. Surface roughness characteristics were also analyzed, but the bright color of the pineapple yellow filament proved to be just as difficult to measure optically as the glitter within the opal green filament, and the results were found to be insignificant.Full Experiment – Mathematical ModelsUsing the results from the first three DOEs, the optimal riser design to increase compressive strength was 50% cubic infill with a shell thickness of 1.4 mm printed in pineapple yellow Prusament PLA. Due to the issues with filament availability and post-processing, the optimum riser design was instead 50% cubic infill with a shell thickness of 0.7 mm printed in opal green Prusament PLA.

Due to the issues with filament availability and post-processing, the optimum riser design was instead 50% cubic infill with a shell thickness of 0.7 mm printed in opal green Prusament PLA. However, as infill percent increased, so too did the mass and print time of the part, resulting in increased costs and lead time which foundries are hoping to reduce. Thus, infill percent was varied from 5% to 50% cubic infill in 5% increments to provide designers with a mathematical relationship between compressive strength and cubic infill percent, allowing them to pick the appropriate infill percent for their strength needs. Parts were also varied in scale (full-scale, three-quarter scale, and half-scale) while maintaining their length-to-diameter ratio. This was done to verify the model for the specified length-to-diameter ratio, as well as to test the potential differences between the scales.

This was done to verify the model for the specified length-to-diameter ratio, as well as to test the potential differences between the scales. All calculations for regression models and linear coefficients of determination (R2) were done using MATLAB.Initially, the polyfit command was used to produce third-order regression models. Across all scales, the third-order models did not produce as close of a fit as hoped due to the presence of outliers. The models were re-run for a first order (linear) regression by removing a single outlier from the full-scale and half-scale data, resulting in a closer fit.

The models were re-run for a first order (linear) regression by removing a single outlier from the full-scale and half-scale data, resulting in a closer fit. The linear regression1861Infillequations and corresponding R2 values are listed in Table 1.Table 1: Linear fit equations and R2 values.Scale Full 3/4 HalfEquation y = 276.812*x - 469.733 y = 174.076*x - 162.200 y = 90.410*x - 160.667R2 0.982 0.982 0.88An example of the compression strength linear regression data for the full-scale riserswith the associated equation is shown in Figure 10.Figure 10 : Full-scale, linear regression for compressive yield strength with outliers removed.Upon re-inspection of the data, the authors found that what were previously thought to be outliers were incorrectly reported measurements, an issue which arose due to the Tinius-Olsen universal testing machine data collection software.

As illustrated by the machine-supplied load versus position plot in Figure 11, some graphs exhibited two distinct linear regions.In certain cases, the software which calculated the 0.2% offset had difficulties determining the appropriate linear region, providing the higher yield point values by mistake. Thus, the compressive strength values resulting from the infill percentages below 20% (at the full- and three-quarter scales) should be ignored for this study.Print Time and Mass EstimationsGiven the material and time constraints associated with this initial case study, alternative infill types were not printed for experimental testing. Instead, print times and masses were1862Figure 11 : Load vs.

Instead, print times and masses were1862Figure 11 : Load vs. Position graph supplied by Tinius-Olsen software highlighting the two linear regions.estimated using the PrusaSlicer software for six different infill patterns (3D patterns: gyroid, cubic, 3D honeycomb; 2D patterns: honeycomb, concentric, and grid) ranging from 5% to 100% in 5% increments. The infill pattern cross sections can be seen in Figure 12.Figure 12 : Infill pattern cross-sections.An example of the plots produced to estimate print time and mass based on scale andinfill pattern is shown in Figure 13 for the full-scale risers.Surface Finish ResultsSurface finish data was collected for wooden risers (uncoated and coated) as well as plastic (Prusament PLA) risers.

Two methods of collection were used: optical and mechanical.1863Figure 13 : Full-scale print time and mass estimations by infill pattern type.Optical methods provided average surface area roughness (Sa) values, whereas mechanical methods provided average linear roughness (Ra) values, due to differences in the measuring techniques. While the resulting outputs are not exactly the same, Sa is an extension of Ra1864and so the values can be compared to one another for rough analysis.Table 2 provides the corresponding values for the wooden and plastic risers. Uncoated and coated wooden risers were provided by Erie Bronze. All wooden riser measurements were completed using mechanical methods. The minimum roughness value of the uncoated wooden risers was 81 µin. Comparatively, the coated wooden risers had a minimum roughness value of 29 µin, a 52 µin difference.

The minimum roughness value of the uncoated wooden risers was 81 µin. Comparatively, the coated wooden risers had a minimum roughness value of 29 µin, a 52 µin difference. It is important to note that the coated risers were utilized for upwards of 20+ years in a green sand foundry, leading to abrasive wear and storage degradation of the coating. Thus, newer coated risers are expected to have lower minimum roughness values.

Thus, newer coated risers are expected to have lower minimum roughness values. Both the coated and uncoated wooden risers had maximum roughness values above 250 µin, the maximum range of the Mahr Pocket Surf roughness gauge for Ra measurements.Table 2: Surface finish results (µin).81 >250 29 >250 207 343 Plastic - Mechanical Min/Max >250Uncoated Wood - Mechanical Coated Wood - MechanicalMin Max Min Max Min MaxPlastic - OpticalRegarding the plastic risers, measurements were attempted both mechanically and optically.Similarly to portions of the uncoated wooden risers, the roughness of the plastic parts surpassed that of the mechanical equipment’s maximum measurement capabilities of 250 µin; thus it was concluded that the surface roughness of all the plastics parts was above 250 µin. Optically, a minimum surface roughness of 207 µin was found, whereas the maximum was 343 µin.

Optically, a minimum surface roughness of 207 µin was found, whereas the maximum was 343 µin. An example of an optical profile (with units of µm) can be seen in Figure 14.Note that the output of the figure is in µm; the data was converted to µin for direct comparison to the results gathered mechanically from the wooden parts. Additional work is needed here to further evaluate the surface roughness of the plastic risers, as there were issues obtaining the surface roughness data optically due to the presence of deep layer lines and glitter present in the filament.However, from the data collected, it can be generally concluded that the uncoated and coated wooden risers are less rough than the plastic risers.

Therefore, post-processing in the form of smoothing or coating the plastic risers will likely be needed for metalcasting moldmaking process implementation purposes.Wood AnalysisThe mean and standard deviation of the compressive yield strength (lbf and psi) were calculated across the three testing scales (full-scale, three-quarter scale, and half-scale). The1865Figure 14 : Example optical profilometry output.average yield strength of the full-scale, uncoated wooden risers (Table 3) was found to be 12,426.67 lbf , which is approximately equivalent to Prusament PLA risers with a cubic infill of about 45%. However, the standard deviations for both the full and three-quarter scales are equal to approximately 100% of the compressive yield strength mean, indicating a high level of variability across the results.The authors believe such variability may be related to grain orientation, moisture content, or knot location among other wood property considerations.

Some wooden risers were also comprised of multiple parts glued together leading to a visual seam (either horizontal or vertical). Thus, seam orientation is another factor to consider since wood is highly orthotropic in properties due to grain structures. It is also possible that in some cases the strength of the glue is being tested rather than the wood itself.Table 3: Wood compressive yield strength data.Scale Avg. YS (lbf ) Std. Dev. (lbf ) Avg. YS (psi) Std. Dev. (psi) Full 3/4 Half12426.6667 8230.0000 563.500013487.0358 8169.4492 24.74871757.3333 2069.0000 319.00001907.5865 2054.0694 14.1421Dimensional TolerancesMinitab was used to evaluate the dimensional accuracy of the wooden parts for both height and diameter. Hypothesis testing was run at an alpha significance level of 0.05.

Hypothesis testing was run at an alpha significance level of 0.05. The expected mean correlated to the modeled height and diameter of the nominal riser at each1866scale.For the wooden risers, only the height at three-quarter scale was found to reject the null hypothesis, indicating that it was not statistically equivalent to the expected mean. All other scales for both diameter and height were statistically equivalent to their corresponding scale’s dimensions. The results for the wooden risers can be found in Table 4.Table 4: Wood dimensional tolerances for height (in). Note: Using the 95% confidenceintervals, LB = Lower Bound, UB = upper bound.Scale Mean 5.006 Full 3.7785 3/4 2.512 HalfStd. Dev. 0.0056 0.0091 0.0099LB 4.992 3.7559 2.4231UB 5.02 3.8011 2.6009Tolerance p-value 0.206 0.032 0.3360.028 0 0.1778Plastic dimensional tolerances can be found in Table 5.

Dev. 0.0056 0.0091 0.0099LB 4.992 3.7559 2.4231UB 5.02 3.8011 2.6009Tolerance p-value 0.206 0.032 0.3360.028 0 0.1778Plastic dimensional tolerances can be found in Table 5. As can be seen by the low p-values, height and diameter are not statistically equivalent to the expected scales’ mean for both full-scale and three-quarter scale plastic risers. At the half-scale, the height was found to be statistically equivalent however the diameter was not. While the plastic risers’ expected means do not fall within the measured lower and upper bounds, their dimensional tolerances (or spread) are much lower than that of the wooden risers. The plastic risers generally displayed a tolerance of a couple thousandths of an inch, whereas the wooden risers showed tolerances between a couple hundredths to a couple tenths of an inch. Thus, the difference of actual dimensions to desired dimensions of the plastic risers can be accounted for ahead of time in order to meet the expected values.

Thus, the difference of actual dimensions to desired dimensions of the plastic risers can be accounted for ahead of time in order to meet the expected values. This should pose no difficulties to foundry designers who are already used to accounting form patternmaker’s shrink when designing foundry patterns.Table 5: Plastic risers dimensional tolerances for height (in). Note: Using the 95%confidence intervals, LB = Lower Bound, UB = upper bound.Scale Mean 3.0065 Full 2.4415 3/4 1.5093 HalfStd. Dev. 0.013 0.1443 0.0103LB 2.9741 2.083 1.4171UB 3.0389 2.8 1.6014Tolerance p-value 0.479 0.148 0.4230.0648 0.72 0.1843Failure ModesThree distinct failure modes were identified for the plastic (Prusament PLA) risers:compression, layer collapse, and brittle crack; each fracture mode is pictured in Figure 15.”Compression” was considered as a uniform progressive failure that occurs as the part yields under load.

Compression failure modes were seen at the full-scale for 40% to 50%, at the three-quarter scale from 40% to 50%, and at the half-scale from 35% to 50% infill densities. ”Layer collapse” indicated a cyclical failure that occurred as layers progressively collapsed under load. Layer collapse failure modes were seen at the full-scale from 30% to1867Figure 15 : Plastic failure modes: compression (left); layer collapse (middle); brittle crack (right).35%, at the three-quarter scale for 15% and 25% to 35%, and at the half-scale from 10% to 30% infill densities. Finally, ”brittle cracking” was a sudden catastrophic failure which occurred due to cracking along layer lines.

Finally, ”brittle cracking” was a sudden catastrophic failure which occurred due to cracking along layer lines. Brittle cracking failure modes were seen at the full-scale from 5% to 25%, at the three-quarter scale from 5% to 10% and at 20%, and at the half-scale at 5% and 20% infill densities.Generally speaking, compressive failure occurs at relatively higher infill densities, brittle cracking occurs at relatively lower infill densities, and layer collapse falls in the middle range of infill densities. Though some exceptions can be noted, e.g. layer collapse at 15% infill for three-quarter scale risers. The authors believe such discrepancies may be due to how the infill is distributed at the lower scales, and consequently where pockets of air at lower infill densities are positioned within the portion of the part receiving the load.Regarding the wooden risers, four failure modes were identified: compression, paintbrushing,splitting, and mixed.

The four failure modes are pictured in Figure 16.1868Figure 16 : Failure mechanisms seen in the wooden risers (both coated and uncoated): (a) compression, (b) paintbrushing, (c) cracking, and (d) mixed.Due to the smaller sample size, generalizations were more difficult to ascertain. However, compression failure modes seemingly related mostly to horizontal seams, whereas split and mixed failure modes mostly related to vertical seams.ConclusionsThree initial DOEs and a full experimental study were conducted to answer the five research questions posed at the beginning of this case study. Regarding each of the research questions, the conclusions are as follows:RQ1: Does infill density (percentage) affect the print time, mass, compressive strength, dimensional stability, or surface roughness of additively manufactured risers of a standard length-to-diameter ratio?– Infill percentage significantly affected print time, height, diameter, mass, and compressive yield strength.

Surface roughness was not found to be significantly affected by infill percentage, but the interaction between infill percentage and infill type significantly affected surface roughness.1869RQ2: Does shell thickness affect the print time, mass, compressive strength, dimensional stability, or surface roughness of additively manufactured risers of a standard length-to- diameter ratio? – Shell thickness did not significantly affect print time, mass, dimensional stability or surface roughness. It did significantly affect the compressive yield strength.RQ3: Does infill type (pattern) affect the print time, mass, compressive strength, dimensional stability, or surface roughness of additively manufactured risers of a standard length-to-diameter ratio?– Infill type significantly affected print time, height, diameter, mass, and compressive yield strength.

It did not significantly affect surface roughness, but its interaction with infill percentage did.RQ4: Does filament color affect the print time, mass, compressive strength, dimensional stability, or surface roughness of additively manufactured risers of a standard length-to- diameter ratio? – Filament color did not significantly affect print time, mass, dimensional stability, or surface roughness of the risers.

Filament color did significantly affect the compressive yield strength of the risers, favoring pineapple yellow.RQ5: What are the print time, mass, compressive strength, and dimensional stability capabilities at various infill densities of additively manufactured risers of a standard length-to-diameter ratio for a constant infill type, shell thickness, layer height, and filament color? – Holding the print type (cubic), shell thickness (0.7 mm), layer height (0.3 mm), and filament color (opal green) constant, the compressive yield strength generally increases with increasing infill percentage. Additionally, the compressive strength increases as the riser scale (full-scale, three-quarter scale, and half-scale) increases.This case study served as a good first step to solving the issues associated with adopting additive manufacturing into the metalcasting patternmaking process.

However, there were limitations to the study and therefore future work is necessary to paint a comprehensive portrait of the overall capabilities of AM in regards to applications for green sand metalcasting. The most logical next steps for this work are to extend the infill percentages past 50% and to consider different colors of Prusament PLA filaments using the same print settings. Additionally, investigation of the effects of other brands of filaments or printers would benefit the research since Prusament and Prusa i3 MK3S series printers are considered as rather high-end and unlikely to be as commonly purchased as less expensive brands of filaments and desktop printers. Comparison of other common infill patterns, such as gyroid, honeycomb, and grid, could be beneficial if their print times fall below and/or strengths significantly exceed that of standard cubic infill.

Extended future work could include an exploration of additional material extrusion filaments such as polyethylene terephthalate glycol (PETG), acrylonitrile butadiene styrene (ABS), polypropylene (PP), and nylon. Carbon fiber-infused1870and other composite-based filaments would also be interesting to investigate.

Carbon fiber-infused1870and other composite-based filaments would also be interesting to investigate. Finally, research involving aging pattern parts would be beneficial to foundries, including the effects of sand abrasion on dimensional accuracy of AM risers, the effects of humidity and absorption of water on the dimensional accuracy and properties of AM risers, and the effects of using aging/opened filament rolls to print pattern parts.DisclaimerThe data collected, analyses performed, and conclusions drawn are those of the authors and may not necessarily represent the opinions of Erie Bronze and Aluminum Company or any of its associated employees from the past, present, or future.AcknowledgmentsThe authors would like to express their most sincere gratitude to several people at Erie Bronze and Aluminum Company for partnering with us on this project: Ms. Amber Evanoff for serving as the main contact throughout the project, Mr.

Amber Evanoff for serving as the main contact throughout the project, Mr. Brad White for the tour of the facility, and Mr. Matthew Firster for answering all of our engineering questions and providing us with wooden pattern parts. Additionally, the following thank yous are due to faculty and staff of Penn State Behrend: Prof. Adam Wielobob for suggesting Erie Bronze as a local foundry for partnership, Mr. Chris Bartlett for machining off the runners from the old wooden risers, and the School of Engineering for access to the Tinius Olsen universal testing machine. Finally, the authors acknowledge Prof. Paul Lynch of the Materials and Manufacturing Research Group at Penn State Behrend for access to the group’s Prusa 3D printers as well as providing the funding for the filament used in this case study.References[1] “AMO—Metal-Casting-August-2022-Public.pdf.” Accessed: Jun. 26, 2023. [Online].

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEvaluating the corrosion performance of wrought and additively manufactured (AM) Invar ® and 17-4PHI.S. Grech*, Dr N.Wint*, Dr S. Mehraban*, Prof. J. Sullivan* and Prof. N. Lavery*Department of materials science, Swansea University, Wales, SA1 8ENAbstractA Renishaw AM 400 was used to produce Laser Powder Bed fusion (LPBF) iron alloy Invar ® and 17-4PH components. Build parameters were systematically changed and the corrosion performance of the samples produced was investigated using a combination of scanning vibrating electrode technique (SVET) and advanced conventional electrochemical techniques. The results indicated that small changes in the density of the LPBF parts due to porosity resulted in large changes to the materials corrosion susceptibility.

The results indicated that small changes in the density of the LPBF parts due to porosity resulted in large changes to the materials corrosion susceptibility. The LPBF samples also demonstrated significantly more variation in pitting potential measurements compared to wrought samples indicating inhomogeneity in the built parts. References to AM samples in this work refer to samples produced using LPBF.IntroductionLPBF is a new and complex technique that prints metallic parts directly from CAD files. This technique has a great number of variables to consider including laser, geometrical and build chamber variables. While there is limited information available regarding the effect of varying build parameters on the mechanical properties of printed components [1]; the way these variables impact the corrosion performance of LPBF parts is largely unknown.

This lack of knowledge regarding the reliability and reproducibility of parts limits the integration of AM into several industries, particularly those in which safety critical components are used. This work aims to identify the differences in the corrosion performance of LPBF and wrought parts and understand the reasons behind these differences.The freedom of design that AM offers is of significant interest to industry including the Defence Support Services sector [2]. The development and increased understanding of AM processes would allow for bespoke parts to be produced as and when required, reducing both the storage space required and the costs of shipping individual components. Consideration must be given to the operating environment and the needs for resistance to shock, vibrations, controlled atmospheres, as well as corrosion, particularly in maritime environments, where components are subject to harsh conditions.

The high cooling rates associated with LPBF could result in different phase ratios, while unmelted powder can result in porous components. Additionally, the effects of anisotropic and heterogeneous microstructure must be understood for AM to be incorporated in industry.503The flexibility of LPBF means that a range of materials can be printed including single and multi-phase alloys. The materials studied in this work were selected with the aim of obtaining as much information regarding the corrosion performance of AM components as possible. To this end, Invar ®, a low thermal expansion alloy comprising of 27% Ni and bal. Fe, was selected alongside 17-4PH stainless steel which has sufficient chromium content to form a passive oxide layer. 17-4 PH stainless steel is vulnerable to pitting corrosion which occurs due to localised breakdown of its passive oxide layer. Passive oxide layers form naturally on some materials surfaces.

17-4 PH stainless steel is vulnerable to pitting corrosion which occurs due to localised breakdown of its passive oxide layer. Passive oxide layers form naturally on some materials surfaces. These layers are usually a few nanometres thick and greatly reduce the rate of corrosion of the metal in the presence of aggressive anions [3]. Localised breakdown of the passive oxide layer can result in the initiation of pitting corrosion. Active pits can perforate the structure of the material and can act as initiation sites for cracks leading to a reduction in the structural integrity of components [4].During this work LPBF was used to produce iron alloy Invar ® and 17-4PH components.

Build parameters are systematically changed and the corrosion performance of the samples produced is investigated using a combination of scanning vibrating electrode technique (SVET) and advanced conventional electrochemical techniques.Experimental MethodsLPBF Parameter Optimisation and Sample Preparation17-4PH and Invar ® cubes with side lengths of 15 mm were printed using a Renishaw AM 400 LPBF system. The build parameters for these cubes were chosen based on an iterative process whereby component density was optimised [5]. Samples were then polished to a 1 μm finish and subject to a range of corrosion tests including SVET and polarisation. For each powder used, build parameters must be optimised to produce high density components with mechanical properties comparable to their wrought counterparts. Information gained from printing in powder of similar elemental composition and geometry can be used as a starting point for this process.

Information gained from printing in powder of similar elemental composition and geometry can be used as a starting point for this process. The volumetric Energy Density (VED) equation is displayed in equation 1,(cid:1848)(cid:1831)(cid:1830) (cid:3404)(cid:1842) (cid:3401) (cid:1831)(cid:1846) (cid:1842)(cid:1830) (cid:3401) (cid:1834)(cid:1845) (cid:3401) (cid:1838)(cid:1846)where P is power, ET is the laser exposure time, PD is the point distance, HS is the hatch spacing and LT is the layer thickness. The VED conveys the quantum of energy which is deposited within each sample during the powder melting process. Orthogonal arrays were created to produce components with variable energy densities an example of which is displayed in table 1. Several repeat arrays were constructed across the build plate to account for any variation in build quality which could arise due to inhomogeneous gas flow or other504(1)geometrical factors.

Several repeat arrays were constructed across the build plate to account for any variation in build quality which could arise due to inhomogeneous gas flow or other504(1)geometrical factors. Density measurements were carried out on 15 mm3 density cubes using an Attention Sigma 700 force tensiometer which measures the change in force recorded as the sample is immersed in deionised water. The software then uses Archimedes principle to calculate the density of each sample. Statistical analysis was then carried out on the data which fed into the iterative process for the optimisation of build parameters. Cubes, optimised for density, were then selected for corrosion testing. Cubes were sectioned such that several samples were attained from each cube whilst removing the cubes outer finish.

Cubes, optimised for density, were then selected for corrosion testing. Cubes were sectioned such that several samples were attained from each cube whilst removing the cubes outer finish. All samples were ground and polished to a 1 μm finish to ensure that any differences in corrosion performance was as a result of the variation of build parameters as opposed to varying surface roughnesses.Table 1- 17-4 PH orthogonnal aray build parametersSample LabelPoint distance (μm)Hatch Spacing (μm)Exposure time (ms)Power (W)Energy Density (J/mm3)1 260 6060 8060 90180 21560 813 460 80100 60120 90250 250100 945 6 7 880 80 100 10080 100 60 80120 60 120 60180 215 215 25067 32 86 3891001009018032Potentiodynamic polarisationPotentiodynamic polarisation allows for information on the corrosion rate, pitting susceptibility and passivity of an electrochemical system to be determined [6].

When performing a DC polarisation scan, the ionic conduction path is provided through the solution separating the working and counter electrodes, while the electrical conduction path is provided through a potentiostat which controls the voltage difference between the working and reference electrode [6]. The potentiodynamic measurements made in this work were carried out using a custom-made plastic sample holder which allowed an electrical connection to be made to the underside of the sample whilst isolated from the electrolyte. This would allow the polished surface of the sample to be immersed and function as the working electrode in this setup. In an505anodic polarization scan the potential is driven in the anodic direction, causing the sample/working electrode to become anodic whilst electrons flow from the sample through the circuit to the counter electrode.

In this work the recorded current density represents the rate at which the anodic reactions are taking place on the sample/working. This type of scanning is appropriate for investigating materials which readily form a passive oxide film such as stainless steels. Cyclic potentiodynamic polarization is a method of potentiodynamic scanning wherein the applied potential is reversed at the highest applied voltage and reduced to below the open circuit potential (OCP). The reverse section of the scan is used to investigate a materials ability to repassivate. Voltages are quoted relative to the saturated calomel electrode (SCE) and a conventional platinum counter electrode was used with the Solatron 1280 potentiostat galvonostat in the standard three cell setup [6].

The scanning direction was reversed in this work when a current density of 0.5 Acm-2 was recorded.Potentiodynamic scanning was used in this work to explore the pitting and repassivation behaviour of a range of 17-4 PH AM samples with a range of energy densities for comparison with wrought 17-4PH samples. The scans were programmed to record the OCP of the samples for 10 minutes before beginning an anodic sweep starting 200 mV below the OCP and continuing to 1.5V. This scanning range was sufficient to polarise the samples to the point of pitting corrosion and complete sample breakdown. Visual inspection of all samples during and post scanning confirmed that pitting corrosion had taken place.SVET ---- For SVET measurements, samples were polished before being immersed in a 0.17M NaCl electrolytic solution. The SVET microtip probe scans in a raster fashion 150 μm from the surface of the sample scanning.

The SVET microtip probe scans in a raster fashion 150 μm from the surface of the sample scanning. The SVET was configured to record data every 0.2 mm, with a dwell time of 50 ms allowing for 5 measurements per point with a probe oscillation frequency of 140 Hz raster scanning across the sample once every hour for 24 hours. This allows the SVET to collect spatial and temporal mechanistic information over a defined area for a set period of time [7]. The SVET registers an alternating potential that is proportional to the electric field strength in the vertical direction. This alternating potential is produced from the oscillation of the SVET tip in the potential field which is generated by the ionic current flux that passes through the electrolyte solution. This point scan data is then stitched together to produce a sample map at each time step with anodic and cathodic regions visible [7].

This point scan data is then stitched together to produce a sample map at each time step with anodic and cathodic regions visible [7]. SVET scanning was used in this work to observe the difference in the corrosion rate between wrought and AM Invar ® while immersed in ambiently aerated 0.17M NaCl electrolyte. Further manipulation of SVET data allows for the mass loss and mass loss rate due to corrosion to be calculated by making use of the Faraday law and assuming that the corrosion activity occurring on the sample remains constant in the interval between scans. A detailed explanation of the mathematical process is provided elsewhere [8].506Results and discussionMaterial Density ProfilesFigures 1 and 2 show the measured density as a function of volumetric energy density of Invar ® and 17-4PH respectively. The plotted values represent average values from multiple repeat readings with the standard deviation of these readings being used for the error bars.

The plotted values represent average values from multiple repeat readings with the standard deviation of these readings being used for the error bars. The curves obtained are a similar shape for both materials and the relationship between density and volumetric energy density is independent of build material. At low energy densities incomplete powder melting results in low measured densities. The curves then plateau and tend towards the natural maximum density value for each material. In Figure 2 the measured density of the samples begins to drop off after ~100 J/mm3. This may be caused by the vapourisation of powder at high energy levels which reduces the homogeneity of melted powder and results in keyholing and build defects [9]. The range of energy density values selected were centred on the maximum density values shown in Figure 1 and Figure 2.

The range of energy density values selected were centred on the maximum density values shown in Figure 1 and Figure 2. Components for corrosion testing were constructed between VED’s of 43.25 – 60.00 J/mm3 for Invar ® and 32.25 - 100.00 J/mm3 for 17-4PH.8 .15;:;- 8.1 E ~ ~ > .-!:= Vl ~ 8 "'O "'O Q) ~ 7 .95 Vl ro Q) ~ 7.98.05::::::I:::::::l::::___1\:::::~:l:t_1:::JY :J_t:1:::1r::::t :::::e::::L:::::f ::::::IIIIIIIIIIIt f' - + - - - - - - - - + - - - - - - - - -+ - I I- -1- -I- - I-1- - I-1- --1- --..,._ - I....

In all cases, the mass lost from wrought samples was less than the mass lost from the AM samples. After 20 hours of SVET scanning the AM samples lost on average 0.3097 ± 0.1606 g.cm-2 of material compared to the wrought value of 0.0032 ± 0.0180 g.cm-2. Figure 3 shows SVET derived false colour maps from scans of a wrought and an AM sample built with a VED of 59.36 J/mm3 and a measured density of (100.1 ± 0.2) % relative to wrought invar.The red regions in the SVET false colour maps represent anodic areas where corrosion has initiated. The difference between the wrought and AM sample is clearly visible with the AM sample displaying several anodic regions forming after 5 hours. Subsequently these anodic regions increase in intensity over time in the top left anodic region from 1.0 A.m-2 after 5 hours to 6.0 A.m-2 after 29 hours. The wrought sample does not display any form of anodic attack within the first 5 hours.

The wrought sample does not display any form of anodic attack within the first 5 hours. After 29 hours there is some low intensity anodic activity along the left boundary of the sample represented by the white region. This low activity does not represent significant corrosion attack. The different scale bars on the wrought and AM sample SVET false colour maps reinforce the difference in anodic activity with the highest level of anodic activity in the wrought sample after 29 hours measuring a current density of 0.3 A.m-2. Visual examination of the samples post immersion confirms the absence of corrosion product on the wrought sample. The corrosion that occurred to the AM sample focussed around the anodic areas as displayed in the optical images in figure 3.

It is possible that build defects on the invar508sample surface allow aggressive halide species to migrate into the sample and corrode the inside of the sample.Wrought sample after 5 hoursAM sample after 5 hours68-7- E6 E -; 5- (.) C C1l 1i5 0 3 24-Current densityAm ·'0.350.150.05E §.Q) (.) C C1l en 05432Current density • 5, Am4320.250002345 6 7890023456Distance (mm)Distance (mm)Wrought sample after 29 hoursAM sample after 29 hours9687 E 6 §.5Q) (.) ~ 4 en 3 i5Current density Am'0.350.150.05E §.Q) (.) C C1l 1i5 i55432Current density~Am6 543220.2500 0234' 567890023456Distance (mm)Distance (mm)Figure 3- SVET derived false colour maps of wrought AM invar with optical images showing samples post corrosion testing509Potentiodynamic ScanningTable 2 displays the mean pitting potentials for tested samples as well as their measured density.

Sample 6, the lowest density sample demonstrates the lowest pitting potential. The fact that the value is negative is not significant as the values are quoted relative to SCE in this case. The data shows that sample 3, the AM sample with the highest density outperformed every other tested sample with the highest measured average pitting potential. However, this sample also exhibited the largest variation in pitting potentials. Further to this the variation of pitting potentials recorded indicates that the wrought samples demonstrate better repeatability than the AM samples. The average variability recorded for the wrought samples are 7% compared to 80% for the AM samples. The larger spread of pitting potentials recorded in repeat measurements of AM samples could be due to sample inhomogeneity with regards to the random location of build defects.

The larger spread of pitting potentials recorded in repeat measurements of AM samples could be due to sample inhomogeneity with regards to the random location of build defects. Any build defect on, or near, the surface in a sample could result in a reduction in the uniformity of the samples passive oxide layer reducing its effectiveness.It is also apparent that a small decrease in the density of AM parts results in a significant reduction in the pitting potential. This could be indicative of samples with densities below a certain threshold being more vulnerable to pitting due to there being more build defects acting as initiation sites for pitting on the surface of the sample.

The difference between the wrought and unpolished values justifies the reasoning behind polishing all samples and discounting surface roughness effects as the polished wrought samples have significantly higher pitting potentials than their unpolished wrought counterparts.Table 2 - Average pitting potential values for AM and wrought 17-4PH samplesSample1 3 4 6 Wrought Unpolished WroughtDensity (%)99.46 ± 0.01 99.78 ± 0.01 99.56 ± 0.01 93.67 ± 0.01 -Average Pitting Potential (V) vs SCE 0.051 ± 0.031 0.332 ± 0.283 0.023 ± 0.017 -0.192 ± 0.190 0.293 ± 0.0090.176 ± 0.021The polarisation curves displayed in Figure 4 convey the difference between the samples with the highest and lowest pitting potentials, samples 3 and 6 respectively. The regions labelled A and B represent the passive region of each sample respectively. The passive region of sample 6 is severely diminished compared to the passive region in sample 3.

The regions labelled A and B represent the passive region of each sample respectively. The passive region of sample 6 is severely diminished compared to the passive region in sample 3. A decreased passive region has been reported previously in literature for AM parts with the510observation being attributed to elemental segregation, retained oxides, and/or porosity [10]. Additionally, the sharp horizontal extensions from the sample 3’s passive region are indicative of metastable pitting where pits are initiated and re-passivated quickly before they can mature in stable pits [11]. The open circuit potential (OCP) is the potential at which the sum of the anodic and cathodic reaction rates on the electrode surface is zero.

The nature of the hysteresis loop of a potentiodynamic scan yields information regarding corrosion performance within the electrolyte. For example, a negative hysteresis loop with a passivation potential well above the OCP indicates that the testing material will not readily corrode while in service in conditions representative of the electrolyte used [12]. The cyclic potentiodynamic polarization loops in figure 5 show that the AM samples passive oxide layer broke down at a higher potential than the wrought sample which is in agreement with the data in Table 2. Although both scans demonstrate a positive hysteresis loop, the wrought sample demonstrates repassivation at the area labelled Repass. (where the reverse scan).

This large difference is the result of disruption of the passive oxide layer at high potentials, with the large hysteresis loop suggesting significant passive film disruption indicating that the material immersed in this solution would be susceptible to pitting over time while in service in the same environment [11].Repassivation of the AM samples is not observed (the cathodic direction scan does not intersect the initial anodic scan) and the hysteresis loop is open. The summary of this data is that the AM sample with the highest pitting potential demonstrates a greater resistance to pitting initially than the wrought sample. However, once pitting has begun and the passive oxide film has been broken down the wrought sample is able to repair its passive oxide layer and repassivate active pits whereas the AM sample cannot and active pits continue to degrade the sample.

This is further demonstrated by much higher current density in the reverse direction of the CPP scans for the AM sample compared to wrought. The reason behind this is not clear and further investigation is needed to understand this.51210°Conclusions and Further WorkThe corrosion performance of wrought and AM Invar ® was compared using SVET. The SVET data showed that anodic regions were formed on the AM sample after immersion for 5 hours whereas the wrought sample showed very little anodic activity. The increased mass loss exhibited by the AM samples compared to the wrought samples demonstrates the reduced performance of AM Invar ® in the marine environment. Further work to investigate Invar ® could include using a micro indenter to create porosity in a wrought Invar ® sample.

Further work to investigate Invar ® could include using a micro indenter to create porosity in a wrought Invar ® sample. This would allow for differentiating reduced corrosion resistance due to porosity or other effects from the LPBF process.The pitting corrosion performance of AM and wrought 17-4PH was investigated using potentiodynamic techniques. The potentiodynamic results indicate that the pitting potential for AM samples varies significantly with small changes in sample density. Whilst the AM sample with the highest measured density, sample 3 demonstrated a higher pitting potential than the wrought samples every other AM sample tested demonstrated a lower pitting potential than their wrought counterparts. This difference was attributed to inhomogeneities in the AM samples. Despite the increased pitting potential of sample 3 the sample was unable to repassivate.

This difference was attributed to inhomogeneities in the AM samples. Despite the increased pitting potential of sample 3 the sample was unable to repassivate. Further work regarding 17-4PH could involve exploring various post build heat treatments in an attempt to reduce the porosity and the non-homogeneity of the AM samples.AcknowledgmentsThe author would like to thank DSTL for funding this research, the MACH1 centre at Swansea university for the use of their facilities and equipment and Dr Natalie Wint, Dr Shahin Mehraban, Professor James Sullivan and Professor Nicholas Lavery for their support and guidance when carrying out this work.513References[1] N. Lavery, J. Cherry, A. Davies, H. M. Davies, S. Browgn and J. Sienz, "Investigation into theeffect of process parameters on microstructural and physical properties of 316L stainless steel parts by selective laser melting.," The International Journal of Advanced Manufacturing Technology, vol. 76, pp. 869-879, 2015.[2] A.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEvaluating the Effect of Z-pinning Parameters on the Mechanical Strength and Toughness of Printed Polymer Composite StructuresBrenin Bales1, Tyler Smith2, Seokpum Kim2, Vlastimil Kunc1,2, Chad Duty1,21University of Tennessee Knoxville, TN2Manufacturing Demonstration Facility Oak Ridge National Laboratory Oak Ridge, TNAbstractTraditional Fused Filament Fabrication methods create a mechanically anisotropic structure that is stronger in the deposition plane than across successive layers. A recently developed pinning process deposits continuous pins in the structure that are orientated in the build direction across multiple layers. Initial studies of this technique have demonstrated the ability to increase inter-layer strength and toughness.

Initial studies of this technique have demonstrated the ability to increase inter-layer strength and toughness. The current study evaluated various z- pinning parameters for carbon fiber reinforced polylactic acid (CF-PLA) structures, including infill percentage, pin length, and deposition pattern. Each of these was found to affect the ability of the z-pin to mechanically bond with the existing lattice structure and had a resulting impact on the mechanical strength and toughness. Initial studies showed an increase in ultimate tensile strength in the Z-axis of around 3.5x. Upon expanding the pinning settings, further studies showed increases of over 35% from the X and Z axis ultimate tensile strength and improved mechanically isotropic behavior.IntroductionFused Filament Fabrication (FFF) is a method of AM in which consecutive layers ofmaterial are deposited onto each other in one direction, typically parallel to the build platform.

The 3D printed structures will be built in the Z-Axis (build direction). Traditional FFF methods create mechanically anisotropic structures due to their consecutive layer printing. When printing layers of material, the strength of the structure is strongest when the beads of material are aligned (0⁰) to one another in the load direction (X-Y Plane), as a byproduct, this creates a very anisotropic structure. Compared to when the beads of material are normal (90⁰) to each other in successive layers, strength falls by 40-85%, but the structure behaves as a quasi-isotropic structure [1–3].When comparing the deposition plane (X/Y) to successive layers in the build direction(Z) the anisotropic behavior of the print leads to a drop in strength. This drop in strength can range from 25-90%. Looking at Table 1, the drop in ultimate tensile strength between ABS materials ranges from 25-79% across both large (BAAM/LSAM) and small (FFF) area additive manufacturing platforms [4–10].

The material of choice for this project is the 15% CF-PLA,508which when comparing the ultimate tensile strength across the X and Z axis, shows a 34% drop in strength[10].MaterialPlatformUTSxUTSzReduction ReferenceABSABS ABS ABS ABS ABSFFF FFF FFF BAAM* LSAM**31 MPa 29 MPa 34 MPa 36 MPa 22 MPa6.6 MPa 14 MPa 18 MPa 27 MPa 11 MPa79% 52% 47% 25% 50%[4] [5] [6] [7] [8]5% Jute-ABS 5% Jute-ABS 13% GF- ABS/LLDPE^ 18% GF- ABS/LLDPE^ 20% GF-ABS 20% CF-ABSFFF FFFFFFFFF BAAM* BAAM*26 MPa 24 MPa39 kg59 kg 67 MPa 67 MPa13 MPa 8.6 MPa1.2 kg11 kg 13 MPa 8.2 MPa50% 64%97%81% 81% 88%[5] [6][9][9] [7] [7]PLAPLA PLAFFF FFF56 MPa 55 MPa24 MPa 37 MPa57% 33%[4] [10]15% CF-PLAFFF53 MPa35 MPa34%[10]Table 1: Mechanical Anisotropy of Extrusion-Based Printed Samples [11]Researchers have come up with many different proposed solutions to alleviate the drop instrength across layers.

Initial work on increasing the inter-layer bond strength has focused on increasing the inter-diffusion of polymer chains across the interface [12]. To increase the inter- layer bond strength it is critical to keep the build temperature to at least the glass transition temperature of the material [12–14].In order to do this, some researchers investigated adding another heat source during the deposition process. A Montana State University research team investigated increasing the interface temperature by introducing a stream of hot air [15]. This method was not effective due to the additional air flow effecting the geometry of the build and not improving the strength of the material. Alternatively, work done at Arizona State University increased the flexural strength by 50% in an ABS (0.03% Carbon Black ABS) by adding a laser- based pre-deposition method to increase the temperature of the deposition surface [16,17].

Oak Ridge National lab also investigated additional heating methods by mounting infrared lamps onto the BAAM printing head, where after preheating, the fracture energy of 20% CF-ABS[18] had doubled. This research further explores a novel printing approach, called z-pinning, that was first presented at SFF in 2017 [19,20].509Z-PinningThe Z-pinning process was developed to increase z-axis strength and reduce mechanicalanisotropy. The Z-pinning process in composite AM manufacturing aims to achieve similar goals to the Z-pinning processes already being used in the traditional composites manufacturing industry to eliminate delamination and improve through-thickness properties [21,22]. The z- pinning process for FFF relies on using the same material throughout the build. The current process uses intentionally created aligned gaps over successive layers (see Fig. 1) to create a void where the pin is dropped through multiple layers.

The current process uses intentionally created aligned gaps over successive layers (see Fig. 1) to create a void where the pin is dropped through multiple layers. The pinning process starts by dropping in the pins through a set amount (set using the pin parameters) of layers (Fig. 1a). The next two steps of the process (Fig. 1b and 1c) show the pin being deposited across multiple layers. As the pins are deposited on top of one another, a “seam line” is created between the two consecutive pins. This seam line is a potential source of a failure. It is important to note that the pins are staggered as to avoid creating a seamline, which can lead to defects in the print (Fig. 1d). Next, after laying down another set of layers, the nozzle will be positioned over the already deposited pin (Fig. 1e and 1f) and stack the pins on top of one another.

1d). Next, after laying down another set of layers, the nozzle will be positioned over the already deposited pin (Fig. 1e and 1f) and stack the pins on top of one another. The z-pinning method’s effectiveness is determined by how well the pins mechanically interlock with the surrounding layered structure[23] and how well the materials bond to one another. This is determined by the various pin parameters that determine the overall shape and size of the pin.Figure 1: The z-pinning process[11]The pin parameters are the pin width, pin length, deposition pattern (seam offset), and the infill percentage. The pins themselves are designed as a prismatic pin with a square cross section with a length that spans a specific number of layers. The pin width defines how wide the pin is, ranging from 1.5mm to 3mm. Pin length determines how many layers the pin will be deposited across, either 8, 12, or 20 layers.

The pin width defines how wide the pin is, ranging from 1.5mm to 3mm. Pin length determines how many layers the pin will be deposited across, either 8, 12, or 20 layers. The deposition pattern of the pin affects where the seams will line up with one another. This is otherwise known as the seam offset. The seam offset is set at510half of the pin length of the pin. This means that the pins will have a staggered seam pattern so to not create an obvious fracture line (See Fig. 2). The final pin parameter is the infill percentage. The infill percentage is defined as the amount of material that is deposited into the pin gap compared to the theoretical volume of the gap itself. how much material is deposited into the void. The current infill percentages being tested are 80%, 100%, and finally an “overflow” at 120%.

how much material is deposited into the void. The current infill percentages being tested are 80%, 100%, and finally an “overflow” at 120%. The ability to “overflow” the pin gap is due to the nature of the rectilinear grid structure part, meaning that the material is laid down in alternating X/Y deposition paths (See Fig. 3).Figure 2: Pin Parameters [4]Figure 3: Rectilinear Grid [11]511Initial studies in z-pinning showed promising results in both improving the z-axis ultimate tensile strength (UTSz) and leading towards quasi-isotropic data in some of the data sets. The pin parameters used for the first round of research consisted of a pin length of 4-8, meaning the pins were 8 layers long and staggered at 4 layers. While the pin width was held at 3W, which sets the width to be 1.5mm.

While the pin width was held at 3W, which sets the width to be 1.5mm. The infill percentages varied from 80%, 100%, 120% with additional data sets at a solid infill (standard rectilinear grid) and 0% infill (no pins in the voids).Figure 4 shows the relationship between the ultimate tensile strength of the X-axis (UTSx) and Z-axis (UTSz) [11]. The pin parameter that showed the greatest ultimate tensile strength was the solid sample set. The solid sample showed a 33% degree of anisotropy between the X and Z axis. Looking at the pinned samples, from 0%-120% samples increased in strength and decreased in anisotropy. The 120% infill samples showed an average UTSz over 3.5x greater than the 0% infills samples (See Fig 4). Additionally, the 120% samples showed the lowest anisotropy.

The 120% infill samples showed an average UTSz over 3.5x greater than the 0% infills samples (See Fig 4). Additionally, the 120% samples showed the lowest anisotropy. Additionally, the goals of the next phase of research is to systematically investigate what effects different pin parameters have on the mechanical strength and isotropy of printed parts.Figure 4: Ultimate Tensile Strength of CF-PLA [11]Experimental MethodsThe printed samples were walls measuring 12.7cm x 1.27cm x 12.7cm (Fig. 5) Printedusing a MakerGear M2 Desktop printer. The material used was 1.75mm diameter CarbonX from 3DXTech. The CarbonX material is a carbon fiber reinforced PLA (CF-PLA) [24] containing a 15% weight carbon fiber reinforced PLA (CF-PLA) [25]. Printing took place at a nozzle temperature of 225⁰C and a bed temperature of 65⁰C.

Printing took place at a nozzle temperature of 225⁰C and a bed temperature of 65⁰C. This slight increase over the recommended512temperatures of 220⁰C (extruder) and 60⁰C (bed) was selected for the additional benefits of increased bonding to the print bed. The software used to generate the g-code was Matlab and was programmed to create a rectilinear grid of alternating X and Y axis deposition paths. The rectilinear grid has a 35% infill pattern with preset voids to insert the pins. The rectilinear grid was deposited with a bead dimension of 0.53 mm wide 0.30mm thick. To create the 35% infill, the beads were spaced apart by 1.87 mm (Fig. 3).m c7 2 1.Figure 5: Print Sample with DimensionsDuring the printing process 4 walls were printed for each pin parameter. This is necessarybecause each wall will yield three tensile samples (for a total of 6 samples), which allows us to have enough samples for a standard deviation, per ASTM D638 [26].

This is necessarybecause each wall will yield three tensile samples (for a total of 6 samples), which allows us to have enough samples for a standard deviation, per ASTM D638 [26] . The walls are then milled to less than 9mm in thickness, which allows the samples to fit inside of the grips of the tensile tester. After milling, the walls were waterjet cut into 3 tensile test specimens per wall. Two of the walls were cut in with perpendicular to the build path (Z-samples), while the other two walls are cut parallel to the build path (X-samples) (Fig. 6). The sample geometry was a modified Type 4 tensile specimen per ASTM D638 (Fig. 7) [26]. The tensile test specimens were a modified geometry due to the need for a sufficient number of pins across the gage area. If there are not enough pins across the gage area, then the pins will have an insufficient effect on the material properties of the samples. Before testing the samples are dried in an oven for at least 8 hours at 50⁰C.

Before testing the samples are dried in an oven for at least 8 hours at 50⁰C. Mechanical testing is conducted using an MTS servo-hydraulic tensile tester equipped with a 110kn load cell with a strain rate of 3 mm/min.513Fig 6: Orientation of Tensile Samples [11]Fig 7: Modified Type 4 Tensile Sample, dimensions in mm [11]This stage of the z-pinning involved expanding the pin parameter set to discover if anyalterations to the pins would give them an advantage in ultimate tensile strength. Additionally, if any of the new pin parameters would lead to a more isotropic data set. The first set of pin parameters that were expanded upon were more 80% infill pins, this is the focus for this study. The set of parameters that were first identified were the 3W (1.5mm) pin width, and the pin lengths of 4-8 and 6-12. After the 3W set, 4W (2mm) and 6W (3mm) were also printed and tested.

The set of parameters that were first identified were the 3W (1.5mm) pin width, and the pin lengths of 4-8 and 6-12. After the 3W set, 4W (2mm) and 6W (3mm) were also printed and tested. Both of those samples were printed at the 6-12 pin length.Results and DiscussionsTensile data will be analyzed in using two methods. First, the strength ratio, which is apercent change between the between the ultimate tensile data of the 0% infill sample (See Fig. 4) and the ultimate tensile strength of the selected sample (See Eq. 1). The strength ratio will allow a direct comparison between how the z-pinning process (and the different parameters) will affect514a unpinned print, specifically with a pin width of 3W[11]. The other metric is a calculation of the anisotropy ratio. The anisotropic ratio is the ratio of UTSz to UTSx for a given sample.

The other metric is a calculation of the anisotropy ratio. The anisotropic ratio is the ratio of UTSz to UTSx for a given sample. This ratio will demonstrate how the different pin parameters effect the individual tensile strengths and how they relate to one another.(𝑈𝑇𝑆0% − 𝑈𝑇𝑆80% 𝑈𝑇𝑆0%) 𝑥 100Equation 1: Strength Ratio ComparisonTensile tests results shown in Figure 8 for all 80% pin widths. The tensile sample with apin width of 3W, and a pin length of 4-8 demonstrated a 53.4% decrease in strength when comparing UTSx to UTSz. The final tensile sample at 3W had a pin length of 6-12. The pin length corresponded to a 63.7% decrease in strength (Fig. 8). The final two sample sets at 80% were a 4W/6-12 and a 6W/4-8. The 4W/6-12 sample was chosen as the “middle ground” in the parameter matrix (Fig 8), which would in theory yield a combination of the best tensile strength and easiest print process.

The 4W/6-12 sample was chosen as the “middle ground” in the parameter matrix (Fig 8), which would in theory yield a combination of the best tensile strength and easiest print process. Looking at Figure 8, the 4W/6-12 sample demonstrated a 52.4% decrease in Ultimate tensile strength, the second lowest. The 6W/4-8 sample was chosen due to the difficulties involved with printing 6W sample sets, it also demonstrated a 48% decrease in strength, the lowest of all the tensile specimens.Ultimate Tensile Strength 80% Samples14.0] a p M12.0[ h t g n e r t S e10.08.0l i s n e T e t a m6.04.0i t l U2.011.65.410.23.78.84.27.53.90.03W 4-83W 6-124W 6-126W 4-8X-AxisZ-AxisFigure 8: Ultimate Tensile Strength of 80% Infill Samples515Table 2 compared the anisotropy ratios of tensile samples according to their pin width and length. The anisotropy ratio is the UTSz divided by UTSx.

The anisotropy ratio is the UTSz divided by UTSx. For the 3W pin widths, the 4-8 sample had an anisotropy ratio of 0.47. While the 3W/6-12 pin length had an anisotropy ratio of 0.36, the lowest of the tensile samples. The 4W/6-12 sample yielded the second highest at 0.48, which combined with the second lowest drop in UTS (Fig 8), validated the testing choice. The 6W/4-8 sample showed the highest anisotropy ratio at 0.52, but only achieved this by decreasing the UTSx when compared to an unpinned 3W sample (Table 3).80% Infill Z/X 4-8 6-12 10-203W 5.4/11.6 = 0.47 3.7/10.2 = 0.364W4.2/8.8 = 0.486W 3.9/7.5 = 0.52Table 2: Anisotropic Ratio of Ultimate Tensile StrengthThe last analysis method used was the strength ratio, described in Equation 1. The strength ratio compared the ultimate tensile strength of the pinned samples, to an unpinned block at 3W pin width, these results are shown in Table 3.

The strength ratio compared the ultimate tensile strength of the pinned samples, to an unpinned block at 3W pin width, these results are shown in Table 3. The first samples at a 3W/3-8 showed a 46% increase in UTSz and an increase of 85% for UTSx (Table 3). The 6-12 pin length sample for 3W only increased the UTSx by 29% and UTSz by 27%. The last two sample sets shown in Table 3 were the 4W/6-12 sample and 6W/4-8. The strength ratio of the 4W/6-12 sample, demonstrated only a 11% increase in the UTSx, but improved the UTSz by 44% (Table 3). When comparing the strength ratios of the 6W/4-8 to the unpinned sample at 3W the UTSx decreased by 5.4%. While UTSz increased by 33.6% (Table 3).Strength Ratio UTSx UTSy3W 6-12 3W 4-8 29% 46% 27% 85% Table 3: Strength Ratio of Ultimate Tensile Strength4W 6-12 11% 44%6W 4-8 -5.4% 33.6%DefectsThroughout the experiment one of the discoveries was how each pin parameter can affectthe printing process of the wall.

For a successful pin creation, the nozzle must deposit the material into the pin void (See Fig. 2b) without coming into major contact with the side walls early. If the pins contact the walls to early, voids will be created (See Fig. 9) [27]. These voids lead to two major sources of error. First, the voids lead to obvious fracture defects that effect the behavior of the sample during the testing process. Additionally, the voids cause the material to pile up before they fall the full length of the pin cavity. This will cause “overflow” errors. Overflow errors result in excess material building up on the printing surface of the build, which then runs into the nozzle during the printing process. Once a significant number of layers has been deposited, the build will sometimes times “step” and fall over.Additionally, analysis of where the defects occurred highlighted how time consuming anddifficult printing at a width of 6W was.

Going forward, the scope of 6W testing will be reduced to focus more on pin parameters that are more practical to investigate.516Fig 9: Voids in Pin Printing[27]Conclusions and Future WorkResearch into z-pinning is still on going, the initial round of research conducted established that the z-pinning method is effective. Prior work [11] concluded that introducing pins into a 35% infill structure was able to increase the toughness and ultimate tensile strength in the z-axis by over 3.5x. Upon expanding the scope of the research to explore a wider range of pin parameters, the effect of the change in pin parameters can be seen directly through the strength ratios of each tensile sample set. The 3W/4-8 tensile sample demonstrated the greatest increase in tensile strength in both axis by 45% when compared to the unpinned sample (See Fig. 4).

The 3W/4-8 tensile sample demonstrated the greatest increase in tensile strength in both axis by 45% when compared to the unpinned sample (See Fig. 4). While pin lengths of 6-12 at both the 3W and 4W pin width only showed modest increases in ultimate tensile strength, 26-29% (3W) and 11-44% (4W). The 6W pin width with a 6-12 pin length proved to be the least advantageous, with a decrease in UTSx by 5.4%, with the second lowest increase in UTSz of 33.6%. Looking at the anisotropy ratios (Table 1), the performance of the 6W/4-8/80% sample showed that while the UTSx decreased, the increased UTS in the Z- axis yielded the best anistropic ratio of 0.52. It is important to note that while the 6W tensile sample led to the best anisotropic ratio, it only achieved this by reducing its UTSx. As more pin parameters are investigated, this drop in strength, and how it affects the anisotropic ratio will be taken into consideration.

As more pin parameters are investigated, this drop in strength, and how it affects the anisotropic ratio will be taken into consideration. Further expansion of the pin parameters into the 100% infill range is planned, to identify the best isotropic ratio and mechanical performance. Additionally, more tests to round out the 80% infill pattern are being conducted, particularly in the 10-20 pin length.517AcknowledgementsResearch sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle, LLC. This work was supported in part by Oak Ridge Institute for Science and Education through the Higher Education Research Experiences Program (HERE). Research was also sponsored by the U.S. Army Combat Capabilities Development Command Aviation & Missile Center.References[1] A. Bellini, S.

ALUATING THE MECHANICAL PROPERTIES WITHIN THE TRANSITION REGION OF MULTI-MATERIAL LARGE-FORMAT EXTRUSION ADDITIVE MANUFACTURINGJ. Brackett1, A. Defilippis2, T. Smith3, A. Hassen3, V. Kunc3, and C. Duty2,31Bredesen Center for Interdisciplinary Research, University of Tennessee – Knoxville, TN 2Mechanical, Aerospace, and Biomedical Engineering, University of Tennessee – Knoxville, TN 3Manufacturing Science Division, Oak Ridge National Laboratory, Oak Ridge, TNAbstractRecent developments in Large-Format Additive Manufacturing (LFAM) have enabled in- situ material changes and production of graded Multi-Material (MM) structures. The Big Area Additive Manufacturing (BAAM) system utilizes a dual-hopper configuration to feed different materials into a screw-based extruder which results in a blended material transition rather than a discrete material boundary, avoiding associated delamination issues.

While this system enables site-specific material deposition at a large scale, the mechanical performance within the transition region needs to be evaluated to determine the impact on end-part performance. Additionally, the transition region can exhibit incomplete mixing, which could further influence mechanical performance. This study investigates the change in mechanical properties throughout the transition region using dynamic mechanical analysis and tensile testing of printed samples. Optical microscopy of the internal microstructure of the printed beads was also utilized to determine the influence of material blending within the transition on mechanical performance.1. IntroductionAdditive Manufacturing (AM) provides a unique ability to construct complex geometries [1, 2], which can be further leveraged with the implementation of multiple materials in the printing process.

Development of Multi-Material Additive Manufacturing (MMAM) has recently seen significant interest [3-8], largely due to the ability to create complex structures exhibiting site- specific properties. MMAM systems must be capable of utilizing more than one material without pre-mixing, pre-compositing, or non-AM post processing treatments [9]. Although there are MMAM systems available for several different types of AM processes, most encounter the same challenge of bonding dissimilar materials during the printing process. These sections of the printed parts are often structural weak points vulnerable to delamination. Vu et al. observed this behavior in T-Peel geometry samples printed with a Stratasys PolyJet Connex 350, noting that fractures occurred more often in the material interfaces. However, aligning interfaces perpendicular to build direction provided increased resistance to fracture at the interface [5, 6].

However, aligning interfaces perpendicular to build direction provided increased resistance to fracture at the interface [5, 6]. Similar behavior has also been observed in Fused Filament Fabrication (FFF) systems. Brischetto et al. observed delamination failure at the interfaces in acrylonitrile butadiene styrene (ABS) and polylactic acid (PLA) sandwich structures [3]. As with the PolyJet system, Kim et al. found that placing interfaces perpendicular to build direction increased resistance compared to interfaces parallel to build direction, but the interfaces remained the most vulnerable to failure [10]. Similarly, Roger and Krawczak demonstrated that placing the interface in multiple z-direction planes increases delamination resistance but still underperforms a single-material counterpart [11].1210One method of mitigating the impact of discrete boundaries between dissimilar materials is the usage of stepwise Functionally Graded Material (FGM) construction.

This method uses discrete steps of intermediate composition between two materials to improve interfacial bonding and create a property gradient. Vu et al. implemented a stepwise FGM core in T-Peel testing and observed an increased resistance to delamination compared to the previous single-material cores [6]. A similar approach was used in the construction of a functional robot shell to vary the elastic modulus from 1 MPa to 1 GPa, providing a balance between flexibility and rigidity that out- performed single-material designs [4]. Some studies have demonstrated the benefits of continuous FGMs, or those with a continuous compositional change and no discrete boundaries, over the stepwise counterparts, especially in bioinspired and biopolymer composite applications [12, 13]. The success of these methods has seen FGMs used in the creation of orthopedic implants [14, 15].

The success of these methods has seen FGMs used in the creation of orthopedic implants [14, 15]. Extensive reviews cover additional unique applications of FGMs using MMAM with a variety of processes and materials [16-18]. Regardless of the situation, FGMs have been used to solve processing issues between otherwise incompatible materials.Recent work on Cincinnati Incorporated’s Big Area Additive Manufacturing (BAAM) system has included the addition of an automated dual-hopper that enables in-situ material changes by rocking the attachment to select which of the two hoppers is currently supplying material to the extrusion system [8]. Initiating a step-change from Material A to B during extrusion creates a graded region of mixed material composition termed the Transition Zone.

Initiating a step-change from Material A to B during extrusion creates a graded region of mixed material composition termed the Transition Zone. As shown in Figure 1, characterization of the progression form Material A to Material B has further defined three stages: an initial Purge Zone of Material A that was already present in the extrusion system followed by the Transition Zone before culminating in a Steady-State extrusion of Material B [19-21]. Further investigation indicated that changes in processing conditions such as rotational speed and screw geometry can affect this transition behavior [21]. In addition to understanding how the material composition changes after switching materials with this system, there is also a need to characterize the mechanical performance.

In addition to understanding how the material composition changes after switching materials with this system, there is also a need to characterize the mechanical performance. This study investigates the change in mechanical response across the transition region and compares it to observed changes in material composition.1211Figure 1: A previously characterized material transition showing the purge (red), transition (yellow) and steady- state (green) zone boundaries marked with semi-transparent overlays.2. ExperimentalTwo structures were printed using the dual-hopper system such that each structure represented a complete material transition: 20 wt % carbon fiber-filled ABS to ABS (CF to ABS) and ABS to 20 wt % carbon fiber-filled ABS (ABS to CF). Dynamic Mechanical Analysis (DMA) and Ultrasonic Assisted Acid Digestion (UAAD) were used to obtain storage moduli and material compositions, respectively.

Dynamic Mechanical Analysis (DMA) and Ultrasonic Assisted Acid Digestion (UAAD) were used to obtain storage moduli and material compositions, respectively. These two properties were then plotted as a function of a normalized extruded volume to compare changes in behavior with material composition.2.1 Sample PrintingAll samples were printed using the BAAM and the novel dual-hopper attachment at ORNL’s Manufacturing Demonstration Facility. Techmer Engineered Solutions supplied two grades of ABS in pelletized form: a neat ABS (HIFILL ABS 1512 3DP) and a 20 wt % CF/ABS (ELECTRAFIL ABS 1501 3DP). Prior to printing, each material was dried at 80 °C for at least four hours. The print used a nozzle diameter of 10 mm (0.4 in) and bead dimensions of 5 mm (0.2 in) by 14 mm (0.55 in). Thermal conditions were kept constant with a bed temperature of 100 °C and a melt temperature of 250 °C.

Thermal conditions were kept constant with a bed temperature of 100 °C and a melt temperature of 250 °C. All samples were printed using a standard mixing screw geometry and screw speed of 300 RPM.Figure 2A shows the programmed print path, which isolated each transition direction to a single structure to simplify analysis. Using a zero percent infill, the print path was intended to create a block U-shaped, single-bead wall with the shown dimensions, which can be seen in Figure 2B. For both the diagram and the image, the red dot represents the point where the dual-hopper switches material and changes transition direction, and the arrows illustrate the travel direction of1212the print head. The final products were 27 layers (approximately 14 mm) tall to provide adequate room for sample creation in both the x- and z-directions.Figure 2: A.) A schematic showing the print path and dimensions.

Each structure shows transition direction; red dots mark both the start and end of a transition print and where material switches. B.) Printed specimens on bed.2.2 Sample MachiningUsing a bandsaw, the completed prints in Figure 2B were first sectioned into 96.5 mm (3.8 in) long plates and then each plate was split in half perpendicular to the print direction to obtain two sides (A and B) of the same composition. This allowed for additional samples from the same composition range, enabling a more thorough and reliable mechanical analysis. As shown in Figure 3A, all plates were milled flat, and four samples in each direction (x and z) per plate were extracted using water jet cutting, resulting in a total of 16 possible samples for a print location. Samples adhered to dimensional specifications found in ASTM D5023 [22]. Due to constraints imposed by the testing equipment (40 mm span), sample thickness was adjusted to remain in tolerance.

Samples adhered to dimensional specifications found in ASTM D5023 [22]. Due to constraints imposed by the testing equipment (40 mm span), sample thickness was adjusted to remain in tolerance. Any samples damaged during the machining process or that did not fall within specified tolerances were discarded. Sample location was tracked using the travel distance of the extruder from the material switch origin (See Figure 2A). For data analysis, sample sets were identified by the travel distance at the center of the x-direction samples, as shown in Figure 3B. distance traveled by the extruder since changing material, and sample sets were defined by the range covered, i.e., the travel distances at each end of the x-direction samples.

distance traveled by the extruder since changing material, and sample sets were defined by the range covered, i.e., the travel distances at each end of the x-direction samples. For data analysis and plotting, the center of these two points was used to identify the location of a set along the print.1213Figure 3: A) Example of the “A” (top) and “B” (bottom) and side plates consisting of the same composition, and B.) A schematic showing how travel distance was identified for each sample set.2.3 DMA Testing ProcedureDMA was conducted using a TA Instruments Discovery Hybrid Rheometer (DHR) in “Three-Point Bend DMA” mode, which mimics a flexural test arrangement. The DHR will perform an operator-programmed frequency sweep and calculate the material response to an increasing rate of deformation while holding constant or varying amplitude and temperature.

The DHR will perform an operator-programmed frequency sweep and calculate the material response to an increasing rate of deformation while holding constant or varying amplitude and temperature. While testing determines the storage modulus (G’), loss modulus (G”), and phase angle (δ), only the storage moduli were reported for the purposes of this study. Focusing on G’ values provided a simplified method for comparing changes in material composition to measured elasticity.The chosen procedure ran a logarithmic sweep over a standard range of 0.1 to 10 Hz, collecting four points per decade. Following ASTM D5023, the strain amplitude was set to 0.01 %, and the initial static force was set at 1 N. Testing was conducted using the largest test frame (40 mm span) compatible with the DHR.

Following ASTM D5023, the strain amplitude was set to 0.01 %, and the initial static force was set at 1 N. Testing was conducted using the largest test frame (40 mm span) compatible with the DHR. The average width and thickness of each sample was manually measured and input into the program prior to test start.For the CF to ABS transition, a significant number of samples were damaged during machining. As a result, some data points were not reported with standard deviations to avoid mis- representing those points as having the required number of tests to report an accurate average. For all other sample sets, at least four samples were tested to obtain an average.2.4 Measuring Material CompositionMaterial composition was obtained at numerous locations in the printed sample, with an increased frequency during transition, to provide context for any observed changes in mechanical response.

Since this study used one neat and one fiber-filled material, composition was tracked through changes in fiber content. Weight percent carbon fiber (wt % CF) was found using Ultrasonic-Assisted Acid Digestion (UAAD), which has proven to be a reliable method for1214accurately characterizing sample composition [19, 20]. As noted, samples were labelled according to distance travelled by the extruder since the material switch. Since the locations of DMA samples did not exactly match that of the available composition data, the measured wt % CF values were interpolated to find the approximate wt % CF of each DMA sample set. The resulting fiber content was used to calculate a corresponding flow rate and improve extruded volume calculations for DMA samples.2.5 Extruded, Residual, and Normalized VolumeBoth material composition and G’ were plotted as a function of volume and compared to determine how changes in material composition affected elasticity.

To accurately calculate the volume extruded, VE, at specific points within the print, the measured volumetric flow rate, Q, was multiplied by the distance traveled by the extruder, d, and divided by known travel speed, ts, as shown in Equation 1.𝑉𝐸 =𝑄 × 𝑑 𝑡𝑠This approach considers any differences between transition directions caused by a change in volumetric flow rate as material changes while also avoiding a reliance on a uniform bead cross- section to calculate volume. To better relate extruded volume to the BAAM, VE is normalized by the standard material volume that occupies the space between the material switch point and the nozzle exit, i.e. the residual volume, VR. With normalized volume, VN, the extruded volume is directly related to the remaining Material A in the system after switching to Material B, which provides clear and meaningful boundaries for observed transition behavior.

All transitions should begin before “1” VN and have roughly equal volumes of Material A on each side of “1”. Transition comparisons can thereby focus on the shape of the curve around VN = 1 and meaningfully represent the dispersion of the discrete material boundary caused by the step-change in material. Furthermore, utilizing both Equations 1 and 2 allows for transitions printed under different condition to be compared.𝑉𝑁 =𝑉𝐸 𝑉𝑅2.5 Plotting against VNBoth carbon fiber content and G’ were plotted as a function of normalized volume. Using a separate y-axis for each data type better visualized the relationship between material composition and G’, which aided in discussing the impact of changing material composition on storage modulus.

Results for z-direction testing were plotted first alone then together with x-direction data to illustrate that observed trends hold regardless of testing direction and then show that the true impact was negligible compared to changes seen in x-direction testing.1215(1)(2)3. Results and Discussion3.1 Effect of Frequency on G’Three-point bend DMA data was collected for four distinct experimental groups: x- and z- directions for both ABS to CF and CF to ABS transitions. Since the programmed frequency sweep calculates G’ at each frequency studied, there was a need to verify that using data from a single frequency would not compromise analysis. Thus, the storage moduli calculated at a high (5.62 Hz) and a low frequency (0.18 Hz) relative to the range covered by the frequency sweep were plotted as two separate trends. The average G’ and standard deviation of ABS to CF x-direction samples for each frequency can be seen in Figure 4 as a function of normalized volume.

The average G’ and standard deviation of ABS to CF x-direction samples for each frequency can be seen in Figure 4 as a function of normalized volume. While the difference between high and low frequency G’ increased in the CF/ABS region, the differences were minimal, and values remained well within the standard deviation. More importantly, the same trend can be seen for both data sets: a plateau in the ABS region followed by a gradual increase during the material change that culminates in a second plateau in the CF/ABS region. This behavior was consistent for the three remaining experimental conditions, validating expectations that data from a single frequency within the LVR could be representative of general trends in mechanical response.

To more closely mimic traditional flexural testing by maximizing the elastic components, high frequency results acted as representative data.Figure 4: X-direction storage moduli at the chosen high and low frequencies against board position for ABS to CF/ABS.12163.2 G’ and Material CompositionFigure 5 shows both G’ and material composition against the normalized volume. As expected, both curves begin prior to and end after VN = 1, and G’ measurements exhibited a lower- value plateau in the ABS region (purge zone) and a higher-value plateau in the CF/ABS region (steady-state zone). In the transition zone, G’ increased with increasing fiber content. Specifically, any change in material composition had an immediate effect on the calculated storage modulus rather than being a lagging or forward indicator.

Specifically, any change in material composition had an immediate effect on the calculated storage modulus rather than being a lagging or forward indicator. The increase in G’ as composition changed across the transition zone would be expected from a mixed material region, indicating that mixing during extrusion was sufficient to prevent degraded material properties. Further, the lack of any visible stair-stepping effect suggests that there were no thresholds in material composition that need to be overcome to see a change in mechanical behavior. As can be seen in Figure 6, z-direction samples exhibited similar behavior as the x-direction samples. This consistent behavior suggests that the transition zone storage modulus can therefore be approximated using a simple rule of mixtures approach based on material composition. Figures 5 and 6 also indicated that standard deviation was influenced by two separate phenomena: carbon fiber content and zone changes.

Figures 5 and 6 also indicated that standard deviation was influenced by two separate phenomena: carbon fiber content and zone changes. For each data set, increasing carbon fiber content led to an increase in standard deviation, and both the onset and end of the transition zone caused spikes in standard deviation. While wt % CF could cause variance through inconsistencies in dispersion and orientation, the spikes when switching zones are most likely due to poor mixing caused by the change in feedstock type and shape.Figure 5: Carbon fiber content and storage modulus plotted against the normalized volume for the x-direction samples from the ABS to CF transition directions.1217Figure 6: Carbon fiber content and storage modulus plotted against the normalized volume for the z-direction samples from the ABS to CF transition directions.3.3 X- vs Z-directionFigure 7 combines the storage modulus data from Figures 5 and 6.

As discussed, the calculated G’ values change with material composition and standard deviation increases with wt % CF regardless of sample orientation. However, the x-direction behavior sees a significant increase in G’ in the CF/ABS region compared to z-direction despite showing similar values while still in the ABS purge zone. One possible cause for this behavior could be limited layer-to-layer boding due to the presence of carbon fibers near the surface of printed beads, which would clearly not be an issue in the fiber-free ABS. Another source could be the preferential alignment of carbon fibers during deposition increasing performance in the x-direction [23].

Another source could be the preferential alignment of carbon fibers during deposition increasing performance in the x-direction [23]. Considering this behavior, the effects of the material transition are best seen through analysis of samples oriented in the x- direction.1218Figure 7: Material composition compared to both x- and z-direction samples from the ABS to CF transition direction plotted against normalized volume.3.4 Effect of Transition DirectionThe direction of material transition was expected to have minimal, if any, impact on mechanical behavior, so Figure 8 shows the composition variation for the x-direction samples for the CF to ABS transition. Like for the ABS to CF direction, G’ shows a direct and immediate relationship to wt % CF for the CF to ABS transition direction. As carbon fiber content decreases, G’ follows suit, creating both expected plateaus and the continuously changing transition region.

As carbon fiber content decreases, G’ follows suit, creating both expected plateaus and the continuously changing transition region. Similarly, standard deviation was dependent on fiber content, decreasing as carbon fiber content decreases throughout the material transition. Once again, standard deviation also exhibited relative spikes at the beginning and end of the material transition, which can again be attributed to inadequate mixing of the differently shaped feedstocks. Continuing to follow trends observed in the opposite transition direction, the transition zone was concentrated near 1 VN. The only noticeable difference in behavior is a slight increase in standard deviation in the ABS region, which is most likely caused by residual carbon fiber stuck on geometry features of the mixing screw design.

It is also worth noting that a few data points during the transition zone suffered from a lack of testable samples and did not have a reportable standard deviation, but they were included as they matched the observed and expected trends.1219Figure 8: CF to ABS material composition and storage moduli for the z-direction samples.4. ConclusionsExtensive three-point bend DMA testing was conducted on multi-material specimens printed with BAAM and the dual-hopper attachment, and both transition directions were analyzed through determination of G’ for x- and z-direction samples. G’ was plotted with previously determined carbon fiber content against the normalized extruded volume to compare changes in mechanical behavior to changes in material composition. For the studied materials, storage modulus exhibited a clear dependence on material composition regardless of sample orientation.

For the studied materials, storage modulus exhibited a clear dependence on material composition regardless of sample orientation. However, x-direction samples experienced a significant change in G’ compared to z-direction specimens, indicating that the effects of the material transition are more apparent for testing of samples oriented in the print direction, which is most likely due to differences in fiber orientation, alignment, and layer-to-layer bonding. Rather than suffering from decreased performance from poor mixing or having to overcome composition thresholds, the multi-material transition zone was characterized by intermediate G’ values, supporting the usage of a transition zone as a replacement for failure-prone discrete boundaries.

Furthermore, consistent behavior could enable the deliberate placement of mixed compositions within printed structures for site-specific properties while also reducing material waste associated with switching materials outside the printed structure. In addition to increasing the average G’, an increase in wt % CF led to an increase in standard deviation. Although the standard deviation in G’ was primarily dependent upon carbon fiber content, it also spiked at the onset and completion of the transition zone relative to surrounding data points. This was attributed to non-uniform mixing when there is a large discrepancy in feedstock percentages.

This was attributed to non-uniform mixing when there is a large discrepancy in feedstock percentages. As expected, these observations held true for the opposite transition direction, confirming that transition direction plays a minimal role in mechanical behavior.Characterization of tensile properties provides an opportunity to further investigate the mechanical properties of the printed material transition region and confirm material composition is the driving force in determining performance. However, there are also opportunities to reduce1220the observed standard deviation through improvements in mixing of the pelletized feedstock during extrusion.AcknowledgementsResearch sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle, LLC.

Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle, LLC. The authors would also like to acknowledge funding from the State of Tennessee and Tennessee Higher Education Commission (THEC) through their support of the Center for Materials Processing. Research supported by the National Science Foundation under Award No. 2055529. Thanks to Cincinnati Incorporated and Techmer PM for provided material and equipment. Further thanks to the UTK MABE Maker Lab.References[1] J. Jiang, A novel fabrication strategy for additive manufacturing processes, Journal ofCleaner Production 272 (2020) 122916. https://doi.org/https://doi.org/10.1016/j.jclepro.2020.122916.[2] J. Jiang, X. Xu, J. Stringer, Support structures for additive manufacturing: A review, Journalof Manufacturing and Materials Processing 2 (4) (2018) 64 Available: https://www.mdpi.com/2504-4494/2/4/64 2504-4494.[3] S.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEvaluating the Surface Finish of A356-T6 Cast Parts from Additively Manufactured Sand MoldsCaitlyn Rodomsky1,2​ and Brett Conner11 ​Youngstown State University 2 ​Applied Systems and Technology Transfer, LLCAbstract:Binder jetting of sand allows for the design and rapid fabrication of complex molds and cores. The surface finish of the printed molds and cores can be transferred to the cast part. A benchmark casting was designed to compare the surface roughness and surface features of several angles on the cast part. The benchmark casting contained surfaces with angles from 5° to 30° at 5° intervals. Benchmark castings from coated and uncoated mold surfaces were evaluated. Angles from 5° to 20° produced a prominent stair step feature.

Benchmark castings from coated and uncoated mold surfaces were evaluated. Angles from 5° to 20° produced a prominent stair step feature. A Keyence microscope was used to measure the arithmetic mean roughness (Ra) and root mean square surface roughness (Rq) of the part surfaces. These measurements are compared to a more conventional contact step features will be discussed. profilometer. The suitability of Ra and Rq for characterizing stairKeywords:surface roughness, stair step features, binder jetting, sand printing, metal castings​1. Introduction. In the [1] United States alone, there are nearly 2,000 metal casting facilities in the country . Although there are many processes for metal casting, sand casting remains one of the most widely used.

Although there are many processes for metal casting, sand casting remains one of the most widely used . Over 70% of all metal castings are produced via sand casting, and the sand casting processes casting industry supplies parts to market sectors such as energy, aerospace and automotiveIn 2016, the global production of metal castings was 104.4 million metric tons [1]​​[2]​​[3]​.​The conventional approach to sand casting tooling is to machine hard tooling such as patterns and core boxes. In many cases, hard tooling can be expensive and have long lead times. This is especially true if the casting is geometrically complex. As shown in , 3D sand printing can produce molds and cores more cost effectively at greater levels of complexity and/or at low production quantities.

As shown in , 3D sand printing can produce molds and cores more cost effectively at greater levels of complexity and/or at low production quantities. 3D sand printing has multiple benefits including on-demand production of molds and cores without the need for hard tooling, core consolidation (thereby eliminating or mitigating core shift, flash, and core assembly damage), affordable low quantity production, and simplifying the fabrication of complex gating designs. 3D sand printing of cores can be combined with conventional hard molds or 3D printed hard molds for greater cost effectiveness[4]​​95. Finally, 3D when exterior mold lines are relatively simple and production quantities are high sand printed molds and cores can be embedded with multi-modality wireless sensor devices to provide data for quality management and solidification model validation[5]​​[6].​​A common method for 3D printing of sand molds is binder jetting.

The ISO standard for additive manufacturing (AM) terminology defines binder jetting as an “additive manufacturing process in which a liquid bonding agent is selectively deposited to join powder materials.” . In binder jetting for sand printing, a layer of sand is evenly spread by a recoater over the build volume. Next, the inkjet head sprays a foundry-ready binder (such as furan or phenolic) anywhere a part intersects the layer. The build box has a false floor that lowers one layer t, in the negative z-direction. This process is repeated until the build is complete. The thickness, build box can be removed from the 3D printer to facilitate depowdering operations.​Like many additive manufacturing processes, layerwise fabrication can produce stair step features in the printed mold or core. The steps in the mold can be transferred to the casting. The stair step features in the casting can change aesthetic appearances.

The steps in the mold can be transferred to the casting. The stair step features in the casting can change aesthetic appearances. They can also require these steps could be stress post-processing machining or grinding for removal. Finally, concentrators that could lead to crack initiation especially in the presence of cyclic loading resulting in fatigue.One can describe stair step features by defining a “print angle.” As shown in Figure 1 on the left, the z-axis is the build direction. The print angle is the angle between the x-y plane in a o given direction (or vector) as shown on the right in Figure 2. Parallel to the x-y plane is =0 θ ​ o ​. In theory, the layerwise =90 θ while perpendicular to the x-y plane (and parallel to the z-axis) is ​ o because layers o printing process should create stair step features at any angle between 0 have a finite thickness,.

As will be seen here, those steps may become less visible at certain t ​ angles because of the background surface topology due to the sand itself.​​​and 90​​​Figure 1 the right, the print angle is shown based on the angle of the surface from the x-y plane.: A schematic of the orientation of the x-y plane and the build direction are shown at the left. On​96[7]​The objective of the project is to determine if the surface roughness influences the mechanical properties of a casting using additively manufactured sand printed molds.

On​96[7]​The objective of the project is to determine if the surface roughness influences the mechanical properties of a casting using additively manufactured sand printed molds. The following research questions will be addressed in this project:1.) What are the characteristics of casting surface features that are associated with the print angle? 2.) Can the surface texture of stair step features be quantified using arithmetic mean roughness (Ra) and the root mean square height (Rq)? Are those measures useful for these and other sand printing features? Can Ra and Rq be expressed as a function of print angle? 3.) Can a refractory coating such as Dura Kote eliminate stair step features? 4.) Are the differences in the surface features between upward facing casting surfaces and downward facing casting surfaces?2.

Methodology2.1 Benchmark casting designThe main purpose of designing a benchmark casting is to use a single casting to determine the size and space of stair step features and to measure the surface roughness of the part surfaces cast at different print angle orientations in the mold. The benchmark casting required multiple surfaces that are at an angle to the build direction. The design for the benchmark casting involved a round four-inch casting that contains six different print orientation angles. These angles consist of 5°, 10°, 15°, 20°, 25°, and 30°. The benchmark is shown in Figure 2.Figure 2: Benchmark casting (left) isometric CAD and (right) the casting itself​The mold was designed to include two benchmark castings per mold: one upward facing and one downward facing. This permitted observation of differences in surface features resulting from upward or downward mold surfaces. Figure 3 shows the gating system for this pair of97castings.

This permitted observation of differences in surface features resulting from upward or downward mold surfaces. Figure 3 shows the gating system for this pair of97castings. The upward facing casting is to the left in Figure 3 and the downward facing to the right. This gating system consisted of one pour cup, one sprue, one sprue basin, two gates to feed the casting, and two tapered feeders/vents.Figure 3: Benchmark part with gating system​The gating geometry shown in Figure 3 prevents a flat parting line. This is not a challenge for 3D sand printing so a staggered parting line that ran at the bottom of the two parts can be seen in Figure 4. This mold contained lifting holds on both sides for the cope and drag. These are inserted to help lift the mold for cleaning and assembly. The pegs on the cope are designed to assist with assembling the cope and drag.

These are inserted to help lift the mold for cleaning and assembly. The pegs on the cope are designed to assist with assembling the cope and drag. The pegs are tapered to prevent rubbing sand-on-sand during insertion of the pegs into the corresponding holes which could distribute debris into the cavity. Assembly instructions were extruded into mold faces and the casting weight of 5.81 lbs was extruded into the top of the cope to ensure that metal was not wasted.Figure 4:Mold for benchmark casting​98Two molds sets were fabricated. The mold design was the same for both molds, but a coating would be applied to the surfaces of one of the mold sets to determine if it improved the surface finish of the cast A356 part. The coating on the sand mold may increase the smoothness of the surface finish of the casting and therefore decrease the post-processing on the cast part.2.2 Flat bar designFlat bars were designed for three point bend fatigue testing.

The dimensions of the flat bar are 101.6 x 50.8 x 6.35 mm (4 x 2 x 0.25 inches). The flat bar was used to maximize the surface area containing stair step features. This also made a useful specimen to study surface features. Fatigue testing results are not included in this paper. A gating system (see Figure 5) was developed to contain multiple flat tensile bars and flat bar fatigue specimens of the same print angle. Four mold sets were fabricated for four print angles: 0°, 5°, 15°, 30°. Several of the fatigue specimens were examined using contact profilometry and comparator plate. 5 and 15° print angle flat bars (both downward facing and upward facing surfaces) are shown in Figure 6. It can be seen that the stairs are more visible on the upward facing surfaces than the downward facing surfaces.Figure 5: The gating system used for the 15° flat bar specimen casting.​99Figure 6: (Far left) 5° flat bar downward facing surface, (Left) 5° flat bar upward facing surface.

(Right) 15° flat bar downward facing surface and (Far right) 15° flat bar upward facing surface​2.3 PrintingThe molds for the castings were produced on YSU’s ExOne S-Max 3D sand printer shown in Figure 7. This is a production-type printer that is used to produce sand molds and cores for the metal casting industry. The printer, including one job box is 6900 x 3520 x 2860 (L x W x 271 x 138.6 x 112.6 inches) and weighs 6,500 kg (14,330 pounds). The build volume is H) mm 1800 x 1000 x 700 (L x W x H) mm (70.9 x 39.4 x 27.6 inch). The build speed of the S-Max is 3 60,000 to 85,000 cm ​/h). The binder system used by the printer was a furan resin which is a furfuryl alcohol. Furan is a foundry resin for green sand. The activator was a toluene sulfonic acid (TSA) and is mixed with the sand just prior to recoating . The mixer is attached to the printer.

Furan is a foundry resin for green sand. The activator was a toluene sulfonic acid (TSA) and is mixed with the sand just prior to recoating . The mixer is attached to the printer. The printer’s recoater spreads a layer of sand that is 0.28 mm thick on the x-y plane of the build box.(​​3 ​/h(2.12 to 3.00 ft​​​​[8]​​Figure 7: YSU’s ExOne S-Max 3D sand printer​100The sand used was IC-80 Oklahoma, which has a grain fineness number (GFN) of 83 which corresponds to an average particle size between 150 to 170 μm. The grain fineness . The higher the GFN the finer the sand and number is the average size of a granular material the smoother the surface of the cast part. The furan binder used was 371381 Enviroset 3D Jet Resin from HA International.[9]​​Two benchmark mold sets were printed. One of the set of benchmark molds was left uncoated. The other was coated with a refractory coating. The coating used was Dura Kote 4482 from HA International LLC.

One of the set of benchmark molds was left uncoated. The other was coated with a refractory coating. The coating used was Dura Kote 4482 from HA International LLC. It is light blue and premixed. Dura Kote is a water-based material coating with a typical density of 19.80 lbs/gal and a Zircon refractory system . The coating was applied manually using a generic paint brush and placed on the print angle surface of the benchmark casting.[10]​​The benchmark castings were cast in alloy A356 at the YSU foundry. No post-casting heat treatment was conducted on the benchmark castings. The flat bars were cast at Lite Metals Company in Ravenna, Ohio. The A356 flat bars were given a T6 temper post-process heat treatment.2.4 MeasurementThe surfaces of the benchmark castings were evaluated using a Keyence Microscope VK-X200 series non-contact 3D laser scanning confocal microscope. The Keyence uses two software packages for scan planning and compiling data: VK-Viewer and VK-Analyzer.

The Keyence uses two software packages for scan planning and compiling data: VK-Viewer and VK-Analyzer. The VK Viewer was used to select the region and direction of scanning. After scanning, the multiple line surface roughness tool in VK Analyzer was used to measure the surface finish of a cast part.A KLA-Tencor D-100 profilometer was used for measurements of the flat bars. The Tencor profilometer is a contact sensor used to measure the roughness of a part. This is unlike the Keyence which is a non-contact system. For both the Keyence and Tencor devices, a measurement length of 12.7 mm was used. The surface finish of the flat bars were also evaluated against a C-9 cast microfinish comparator plate as shown in Figure 8. This allows for visual and tactile evaluation of the surface and provide a surface roughness in Imperial units.

This allows for visual and tactile evaluation of the surface and provide a surface roughness in Imperial units. The comparator plate is a subjective method of measurement.101Figure 8:A cast microfinish comparator plate used to evaluate surface finish of the flat bar​3. Results and discussion3.1 Benchmark casting​q were taken on all six angles on each of the benchmark castings (Up/Bare Mold, Down/Bare Mold, Up/Coated Mold, Down/Coated Mold) using the Keyence Microscope multiple surface roughness line tool. The scans were in the center of each of the print orientation angles and taken down the steps in a 12.7 mm length. Two scans were taken per print orientation angle and then averaged to formulate the Ra and Rq values.The measurements of R​a and RFigure 9: Up/Bare Mold Benchmark Casting​102Figure 10:Down/Bare Mold Benchmark Casting​In both Figure 9 and Figure 10, the stair steps are more visually prominent at angles less than or equal to 15°.

By 25° or 30°, the stair steps cannot be distinguished from the background surface geometry from the sand itself. Besides stair steps, other surface defects can be seen. For example, there are depressed defects on the bottom left hand corner of 5° and 20° angle surfaces from the Up/Bare Mold benchmark casting. the print orientation angles. These can be seen in the bottom left hand corner of the 15° and 20° angles. There is also an elevated defect at the bottom middle of the 25° angle of the Up/Bare Mold benchmark casting. These could be caused by the cleanout of the molds, the packaging, or during transportation. Line-type indents can be seen on the 20°, 25° and 30° angles from the Down/Bare Mold benchmark castings. These can be caused by heavy brush strokes during sand clean out.Figure 11:Up/Coated Mold Results​103Figure 12:Down/Coated Mold Results​Figures 11 and 12 show the surfaces of the castings from the coated molds.

Stair step features can still be seen with the naked eye for the shallow print angles even up to 20° in Figure 12. It is apparent that the Dura Kote coating does not remove the step features. The heavy application of the coating by brush stroke for the shallow angles changes the surface texture with additional rises and depressions. The coating was also not evenly distributed throughout the entire angled surface. Keyence scans were only taken from the center of the part, so the edges were not included in the measurements.Closer examination of the measurements for the bare mold benchmark castings show that the Ra and Rq are higher on the upward facing surfaces than the downward surfaces. As shown for the bare mold castings in Figures 13 and 14, this difference between upward and downward surfaces is consistent regardless of print angle. It is possible this has to do with the orientation of the molds during printing. As shown in Figure 4, the downward facing casting mold faced upward.

It is possible this has to do with the orientation of the molds during printing. As shown in Figure 4, the downward facing casting mold faced upward. Unbound sand would sit on top of the mold, which would compress the stair steps. This loose sand could also abraid the stair step features during removal. The surface roughness on the downward facing surface may also be affected by gravitational effects during the casting pour.Ra and Rq for the castings made with coated molds can be shown in Figures 15 and 16. For the coated molds, this higher Ra and Rq on upward facing surfaces holds true for only 5° to 20° print angle surfaces.

For the coated molds, this higher Ra and Rq on upward facing surfaces holds true for only 5° to 20° print angle surfaces. The 25° and 30° surfaces show the opposite with higher Ra and Rq for the bottom surfaces.104Figure 13:Figure 14:Ra for upward and downward facing surfaces for the benchmark casting with the bare mold.​Rq for upward and downward facing surfaces for the benchmark casting with the bare mold.​105Figure 15:Figure 16:Ra for upward and downward facing surfaces for the benchmark casting with the coated mold​Rq for upward and downward facing surfaces for the benchmark casting with the coated mold.​106.3.2 Flat barsThe Tencor contact profilometer measurements include 0° surfaces which would show a surface without any steps in it and only a sand surface texture. As shown in Figures 17 and 18, the Ra and Rq values of the top and bottom surfaces for the 0° and 30° print angles are very close.

As shown in Figures 17 and 18, the Ra and Rq values of the top and bottom surfaces for the 0° and 30° print angles are very close. However, the 5° and 15° print angle surfaces show the same difference between the upward surface Ra and Rq values versus the bottom ones.Figure 17:Ra for upward and downward facing surfaces for the flat bar castings​107Figure 18:Ra for upward and downward facing surfaces for the flat bar castings​The comparator plate was then used to evaluate the surface roughness of the flat bar specimens. Both visual inspection and touch was used. The 0° plate was evaluated to be most similar to the 420 RMS standard on the comparator plate. 420 RMS would be considered at the high end of the RMS expected for permanent mold castings (200-420) and in the middle of the range of normal non-ferrous green sand (300-560). Next, the 5° and 15° plates were evaluated to be most similar to the 720 RMS standard on the comparator plate.

Next, the 5° and 15° plates were evaluated to be most similar to the 720 RMS standard on the comparator plate. 560 RMS would be on the highest end of the range for normal non-ferrous green sand (300-560) and at the lowest end of the normal ferrous green sand (560-900). Finally, 30° the plate was evaluated to be most similar to the 560 RMS standard on the comparator plate. 720 RMS would be in the middle of the range for normal ferrous green sand.3.3 Analysis of idealized surfacesAs observed in the benchmark castings, the stair steps are more visually prominent at angles less than or equal to 15°. At greater angles the steps become less visible at greater angles, and the surface of the casting appears more like the surface expected from green sand. One way to examine this is to calculate Ra and Rq for an idealized set of stair steps for surfaces at various print angles.

One way to examine this is to calculate Ra and Rq for an idealized set of stair steps for surfaces at various print angles. This would then be compared to Ra and Rq calculated for a idealized sand topology: a series of close packed spheres.108We will start with the stair steps. As a reminder, the arithmetic mean roughness (Ra) indicates the average of the absolute value of height along the reference length (in our case, 12.7 mm). We will calculate in 20 steps along the reference length.R = n alr R = 1 lr ∫ 0 um(abs Z i = t Z mean = aaZ| n| dx (1)1 * si − Z mean otal height verage height(Z) ) (2)N = 2 0Root mean square height (Rq) – indicates the root mean square along the reference lengthR = √ 1qN N ∑ n=1Z 2 n(3)Figure 19 is an image of one stair step to show the distance of one stair step per angle.Figure 19:A graphical representation of a single stair step feature​, to determine the distance of one stair step per angle.

d Solve for the distance, ​​0.28mm ds inθ=(4)0.28 d = sinθ θ is used in this project relates to the print angle, which consisted of 5°, 10°, 15°, 20°, 25°, and 30°. The distance for one stair step at the given angle are below:5° has a one stair step distance of 3.2126 mm 10° has a one stair step distance of 1.6125 mm 15° has a one stair step distance of 1.0818 mm 20° has a one stair step distance of 0.8190 mm 25° has a one stair step distance of 0.6625 mm 30° has a one stair step distance of 0.5600 mmRa and Rq height were then calculated. The maximum allowed distance of the line profile given by the Keyence measuring tool is 0.5 inches or 12.7 mm. This is the distance that was used109to determine the number of stairs per given angle of length 12.7 mm.We now show the calculation process for Ra and Rq for the 5° print angle.

This is the distance that was used109to determine the number of stairs per given angle of length 12.7 mm.We now show the calculation process for Ra and Rq for the 5° print angle. The otherangles follow the same approach.Number of Stairs12.7 mm = 3= 3.2126.9683 stairs (5)Figure 20:Determining distance of a stair step​The number of steps per angle in a 12.7 mm length are shown in Table 1.Table 1:Number of steps per given angle​h Next, reference length of 12.7 mm. s in5° c os5°landare found to determine the equations of the piecewise function over the given​​l 0.28 mm h 0.28 mm= =→ l → h= 0 = 0.0244 (6) .2789 (7)Equations of the piecewise function are below. There are eight equations used to make up the piecewise function.

To calculate the Ra and Rq value, a .1395 mm this case,was calculated based on theZ mean = 0.Thus Ra and Rq values are as follows for 5°:​R = 7 R = 8a q0.70 1.19The same procedure was carried out for all remaining angles (10°, 15°, 20°, 25°, and 30°)to achieve the Ra and Rq values, as seen in Table 25°10°15°20°25°30°Ra70.7068.9565.5863.4263.5861.35Rq81.1980.0176.1173.3873.2075.84Table 2:Calculated values of Ra and Rq for idealized stair step features at given angles​From the literature, binder jetting parts can have green state densities between 44% and 57% . We now consider an idealized model where the sand particles are taken as spheres. Since the packing factor of a simple cubic structure is 50%, we will model the packed sand as an array in a simple cubic structure. A rough direction can be posited as the direction diagonally across the unit cell as shown in Figure 21.

A rough direction can be posited as the direction diagonally across the unit cell as shown in Figure 21. The roughness calculation will be taken across the particles and gaps in this direction for a length of 12.7 mm. Figure 22 is a graphical visualization of the array of idealized sand particles along this orientation.[11]​​111.Figure 21:A left is a graphical representation of a simple cubic array of idealized sand particles. The​arrow points in the direction that will be used in the calculationsFigure 22:A depiction of idealized sand particles with gaps between them​Based on the the GFN of the sand used in the printing, the average particle diameter ranges from 150 μm to 170 μm.

The same methodology is employed but with a piecewise height function of:For sand particles:For gaps:| = √ Z| D| | = | Z| | | 2D 2 22 + x(20)​(19)where increased to for 150 μm, 160 μm, and 170 μm are shown in Table 3.D is the idealized spherical sand particle diameter. The number of steps used was =40 because of the small size of the particles. The calculated values of Ra and Rq N ​​​150μm​160μm​170μm​Ra61.6167.6071.22Rq64.2870.0474.06Table 3:Calculated values of Ra and Rq for idealized close packed, spherical sand particles​The results from both the ideal stair step analysis and the ideal particle analysis are plotted as a function of print angle. Ra is shown in Figure 23. The average roughness of the idealized stair steps decreases as the angle increases. Ra for the idealized sand particles is not a function of print angle and is therefore constant.

The average roughness of the idealized stair steps decreases as the angle increases. Ra for the idealized sand particles is not a function of print angle and is therefore constant. For all angles, Ra for the stair steps is less than those for idealized sand particles with a diameter of 170 μm. That diameter that is the high end of the particle size distribution. The Ra from the stair steps intersects that of the 160 μm idealized sand particles between 10 and 15°. The intersection for 150 μm is near 30°. This analysis appears112to correlate with the visual observation of stair steps disappearing into the background surface texture from the sand at angles greater than 15°.Rq is shown in Figure 24. Rq from the stair steps does not reach the Rq values 150 μm or 160 μm at any angle. However, the stair step Rq does meet the value of Rq 170 μm around 20°.

Rq from the stair steps does not reach the Rq values 150 μm or 160 μm at any angle. However, the stair step Rq does meet the value of Rq 170 μm around 20°. This may mean that Ra is more relevant to the visual appearance of the steps than Rq.It should be noted that the calculated values of Ra and Rq for the idealized simple cubic sand particle array was much greater (2 to 2.5X) than the measured values for the most representative print angle: 0° (it would have no steps). This is not surprising in that real sand is not packed in a simple cubic array. Real sand is also not spherical in shape.

This is not surprising in that real sand is not packed in a simple cubic array. Real sand is also not spherical in shape. The solidification of the casting and the resulting shrinkage would also be determining factors in smoothing the roughness of the casting.Figure 23:Calculated average roughness, Ra, plotted against print angle for the idealized stair step​features and the idealized sand particles.113Figure 24:Calculated root mean squared roughness, Rq, plotted against print angle for the idealized stair​step features and the idealized sand particles.4. ConclusionsSeveral conclusions can be made from the observations above. First, stair step features are visually prominent on the cast surfaces at low values of print angle less than or equal to 15°. At higher angles they become visually indistinguishable from the background texture of the bound sand particles.

At higher angles they become visually indistinguishable from the background texture of the bound sand particles. Using the analytical calculations for Ra, it can be shown that at low print angles, the average roughness of a set of idealized stair step features will be greater than that for a idealized set of a simple cubic packed array of spherical sand particles with particle diameters of 150 μm or 160 μm.It would be desirable to have a mathematical function for Ra and Rq that is dependent on the print angle (i.e. Ra( )). While it is possible to create a function dependent on print θ ​ angle for Ra and Rq based on the idealized analysis, the variability in the measured surface roughness does not readily support such a function based on measurement.) and Rq ( θ ​​​For bare molds, upward facing surfaces on the castings had more visually prominent steps and higher Ra / Rq than downward facing surfaces.

The possible reasons include: abrasive flow of molten metal onto upward mold steps facing downward casting surfaces, the cleaning114practices used to remove loose sand from molds, and the spreading of the furan binder from the stair step risers into the unbound sand causing a fillet in the mold (an observation noted inThe Dura Kote refractory coating did not erase the steps. The stair steps were still visible. This does not exclude the potential of other refractory coatings eliminating these steps. In most cases, the measured values of Ra and Rq for the castings with coated molds were generally less than those for bare molds.Looking ahead, there are several areas still requiring research. First, there is a need to determine the effect of stair step features on static tensile and cyclic fatigue properties of a casting. It would be desirable to determine if a relationship exists between mechanical properties and the print angle.

It would be desirable to determine if a relationship exists between mechanical properties and the print angle. Second, more coatings need to be evaluated to determine if they can remove stair steps features thus eliminating subtractive post-processing operations specific to those features. Finally, it is also desirable to determine if the stair step features affect the performance of fluorescent penetrant inspection (FPI). In particular, it is important to answer two questions: 1.) Do the stair steps appear as linear indications? 2.) Does the background caused by the stair steps features hide actual linear indications?ACKNOWLEDGEMENT This effort was performed through the National Center for Defense Manufacturing and Machining under the America Makes Program entitled “Maturation of Advanced Manufacturing for Low Cost Sustainment (MAMLS)” and is based on research sponsored by Air Force Research Laboratory under agreement number FA8650-16-2-5700. The U.S.

The U.S. Government is authorized to reproduce and distribute reprints for Governmental purposes notwithstanding any copyright notation thereon. This paper has been approved for public release with case number 88ABW-2018-3834.DISCLAIMER The views and conclusions contained in this document are those of the authors and should not be interpreted as necessarily representing the official policies, either expressed or implied, of the Government.References1. Census of world casting production: global casting production growth stalls. Modern Casting. Dec 2017: 24–28.2. Banchhor R, Ganguly SK. Optimization in green sand casting process for efficient, economical and quality casting. Int J Adv Engg Tech/Vol V/Issue I/Jan -March. 2014;25:115[8]​).​29.3. United States. Environmental Protection Agency. Office of Compliance. EPA Office of Compliance Sector Notebook Project: Profile of the metal casting industry.

aluation of 3D Printable Sustainable CompositesDavid A. Roberson 1,2,*, Carmen R. Rocha 1,2 , Monica Piñon 1,31. Polymer Extrusion Lab, W. M. Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, TX 79968 2. Department of Metallurgical and Materials Engineering, The University of Texas at El Paso, El Paso, TX 799683. Department of Mechanical Engineering,The University of Texas at El Paso, El Paso, TX 79968Corresponding Author: D. A. Roberson, droberson@utep.eduAbstractPolylactic acid (PLA) is rapidly becoming the mainstay material for use in desktop grade 3D printers based on FDM technology in part due to the environmental sustainability of this polymer. While biodegradability is an advantage; as compared to other materials used by FDM- type platforms, there is a lack of desirable physical attributes. The work presented here evaluates the altering of the physical properties of PLA through the addition of sustainable additives.

The work presented here evaluates the altering of the physical properties of PLA through the addition of sustainable additives. Here, the physical properties of PLA were modified while, at the same time the two desirable aspects of 3D printer compatibility and biodegradability were retained. Rheological analysis of the material systems was performed by dynamic mechanical analysis and failure analysis of 3D printed tensile specimens was carried out through the use of scanning electron microscopy. Finally, biodegradability of the novel PLA-based material systems was assessed based on in-soil exposure testing.Key Words: 3D Printing; Materials Characterization; Polymer Matrix CompositesIntroductionDesktop-grade additive manufacturing (AM) platforms based on Fused Deposition Modeling (FDM) technology have become a mainstay among home users, educators, and entrepreneurs.

Predominantly relying on acrylonitrile butadiene styrene (ABS) or polylactic acid (PLA) as feedstock, there has been a recent trend of PLA-specific platforms such as the fifth generation MakerBot Replicator and the Printrbot Simple 3D printer. The reason for this shift in material type to PLA is debatable and could be related to perceived health detriments due to the use of ABS in 3D printing applications [1] or the environmentally friendly nature of PLA due to its biodegradability [2, 3].PLA is a biopolymer derived from renewable resources such as corn and sugarcane and while there are examples of other polymers; namely polyethylene, which have been synthesized from the same sustainable resources, the biodegradable aspect of PLA sets it apart from other engineering plastics.

The need for biodegradable polymers stems from the environmental impact of the use of polymeric waste most exemplified by the great garbage patches found in both the Atlantic and Pacific oceans [4-6].914Though PLA does have the unique aspect of being a biodegradable biopolymer; other than 3D printing, the main use of this material is food packaging and disposable utensils. To find a broader range of applicability, the physical attributes of PLA must be altered. As demonstrated by Roberson et. al, [7] expanding the applicability of FDM-type 3D printing also lies in the altering of the material properties of feedstock constituents. Therefore, the broadening of the physical capabilities of PLA will contribute to expanding the applicability of polymer extrusion 3D printing technologies.Several instances of the development of 3D-printer compatible polymer matrix composites (PMC)s with specific end use applications can be found in literature. For example, Shemelya et al.

For example, Shemelya et al. [8] demonstrated a tunable polycarbonate matrix tungsten composite for 3D-printable radiation shielding while Masood and Song [9] demonstrated the capability to alter the heat transfer kinetics of nylon by compounding either copper or iron particles where the end goal was the creation of 3D-printable injection mold tooling.The work presented in this paper demonstrates the altering of the physical properties of PLA while maintaining 3D printer compatibility and biodegradability. Polymer matrix composites were created from a PLA base resin where the filler materials were sustainable substances. Here, we combined PLA with NaCl, jute plant fiber and organo nanoclay pigments. Sodium chloride was chosen due to its abundance, while jute plant fiber was chosen due to examples in literature if the use of this constituent in plant fiber reinforced PMCs [10-13].

Sodium chloride was chosen due to its abundance, while jute plant fiber was chosen due to examples in literature if the use of this constituent in plant fiber reinforced PMCs [10-13]. Palygorskite-based (palygorskite is an organo nanoclay) pigments, originally developed by the Mayan Culture of Mesoamerica were chosen for this study based on work demonstrating an improvement to the mechanical properties of high density polyethylene (HDPE) [14].Experimental ProcedureFor this work, Ingeo PLA, Grade 4043D (NatureWorks, LLC, Minnetonka, MN, USA) was extruded into a polymeric monofilament through the use of a twin screw extruder/compounder (Dr. Collin Model ZK 25T, Dr. Collin, Ebersberg, Germany) equipped with intermeshing co- rotating screws. Baseline PLA and composite PLA-based filaments were extruded to a target diameter of 1.75mm, one of the two standard diameters common among material extrusion 3D printers (the other being 3mm).

Baseline PLA and composite PLA-based filaments were extruded to a target diameter of 1.75mm, one of the two standard diameters common among material extrusion 3D printers (the other being 3mm). In the case of composites, a masterbatch of concentrated material was first created by manually kneading PLA pellets with filler materials through the use of a Carver hot press (Carver, Inc., Wabash, IN, USA). The masterbatch was then granulated with a Brabender GranuGrinder (C.W. Brabender South Hackensack, NJ, USA) and subsequently combined with virgin PLA material in the extruder system.Test specimens were fabricated according the Type V dimensions as described in the ASTM D638-10 standard by printing test specimens with a MakerBot Replicator 2X (MakerBot Industries, Brooklyn, NY, USA) or a Rostock Max Delta 3D printer (SeeMe CNC, Goshen, IN, USA) where the latter unit was modified with an E3D V6 hotend (E3D-Online Limited, Chalgrove Oxfordshire , UK).

Tensile testing was carried out with an Instron® 5866 (Instron, Norwood, MA) tensile testing machine equipped with a 10kN load cell. Dynamic mechanical analysis (DMA) was also performed on all material systems with a TA Q800 (TA Instruments, New Castle, DE, USA).915Four individual filler materials were combined with PLA in the formation of sustainable 3D printable polymer matrix composites: 1) Sodium chloride (NaCl) (Sigma); 2) jute plant fiber sourced from rope purchased from Home Depot; 3) MayaCrom® Blue; and 4) MayaCrom® yellow ( both from Mayan Pigments. Inc., El Paso, TX, USA). The filler materials evaluated in this study were chosen based on the perceived environmental sustainability.

Inc., El Paso, TX, USA). The filler materials evaluated in this study were chosen based on the perceived environmental sustainability. All additives were blended at a 5% (by weight) ratio to PLA.Fracture surface analysis was carried out through the use of a tabletop scanning electron microscope (SEM) operating with an acceleration voltage of 15kV and equipped with a backscatter electron (BSE) detector (Hitachi, Model TM-1000,Hitachi High-Technologies Europe GmbH, Germany). To reduce charge effects within the SEM, all samples were sputter coated with a gold palladium target for 30s with a Gatan Model 682 Precision Etching Coating System (Gatan, Inc., Pleasanton, CA, USA) prior to microscopic analysis.ResultsTensile test results (Fig. 1) revealed that the addition of jute plant fiber and MayaCrom® Blue led to tensile test specimens with lower ultimate tensile strength values lower than those printed from PLA alone.

1) revealed that the addition of jute plant fiber and MayaCrom® Blue led to tensile test specimens with lower ultimate tensile strength values lower than those printed from PLA alone. This decrease in strength may be due to the degradation of the organic components of each additive as noted in Torrado et al. [15] However, the addition of jute plant fiber increased the % elongation specimens were able to endure during the tensile test process. The addition of NaCl did not greatly the mechanical properties.affectSubjecting the material types testing to DMA evaluated here the differences revealed thermomechanical behavior of the composite materials as compared to the baseline PLA resin.

The addition of both NaCl and jute plant fiber to PLA reduced slightly the onset of glassy behavior as determined by the storage modulus curve, however the temperature at which the maximum tan δ occurred was increased, which effectively broadened the thermal window of elastic behavior as compared to baseline PLA (Table 1.) where the addition of jute increased this window by ~8.4°C. In the case of NaCl, though the temperature at which the maximum tan δ occurred was increased, the actual value of this parameter was virtually the same as PLA (1.8) meaning the elasticity was not altered. The addition of both types of MayaCrom®inFig. 1. Tensile test results of materials tested here.916pigments as well as jute plant fiber decreased the maximum tan δ indicating and increase in the elasticity of these composites as compared to the PLA base resin.Table 1.

Fracture surface analysis revealed NaCl to adhere well to the PLA matrix. Crack propagation is evident on the surfaces (highlighted in Fig 3) of the individual particles. There is also no evidence of particle pull out nor were there any gaps between the particles and the matrix. By contrast the fracture surface of the PLA/jute plant fiber composite possessed features indicating a poor interface between the fibers and the matrix (indicated by white arrows in Fig. 1b). The addition of MayaCrom® pigments had the effect of inducing voids within the PLA matrix which accounts for the decrease in tensile strength as compared to baseline samples.theFig. 2. SEM Micrographs of materials tested in this study: a) PLA; b) PLA/jute plant fiber; c) PLA NaCl; and d) PLA/MayaCrom® Blue.917Of the additives tested here, NaCl produced deeper investigation.

2. SEM Micrographs of materials tested in this study: a) PLA; b) PLA/jute plant fiber; c) PLA NaCl; and d) PLA/MayaCrom® Blue.917Of the additives tested here, NaCl produced deeper investigation. While the addition of jute plant fiber increased the glass transition window, the poor adhesion to the matrix at room temperature was a detractor to the overall performance of this exhibited characteristics indicating a robust bond with the PLA matrix, another experiment was carried out where jute plant fiber was first functionalized in to an compounding. The result was a decrease in % elongation while the UTS remained the same (Fig. 4). Further analysis of the fracture surface via SEM (Fig. 5) revealed the presence of polymer fibrils joining the matrix to the jute fiber which formed during the tensile testing process.

4). Further analysis of the fracture surface via SEM (Fig. 5) revealed the presence of polymer fibrils joining the matrix to the jute fiber which formed during the tensile testing process. The presence of these fibrils indicated a robust bond had formed between the matrix and the fiber as the formation of fibrils is cause by the elongating of polymer chains under tension. The decrease in % elongation in conjunction with yield strength values similar to those of the non-functionalized samples indicates an increase in stiffness caused by an increase in the bond strength between the fiber and the matrix.results worthyofcomposite. Since NaClaqueous NaClsolutionpriorFig. 3 NaCl adhered well to the matrix and exhibited a robust interface and shared crack propagation features with the matrix.Fig.

Since NaClaqueous NaClsolutionpriorFig. 3 NaCl adhered well to the matrix and exhibited a robust interface and shared crack propagation features with the matrix.Fig. 4 The effect of NaCl functionalization of jute plant fiber on the mechanical properties of PLA/jute fiber composites.Further analysis of the addition of NaCl to PLA on 3D printed tensile specimens entailed examining the effect of NaCl on the to raster mechanical property sensitivity orientation of 3D printed test specimens. Here two raster patterns were used to print the tensile test specimens: 1) a layers crosshatched raster pattern where alternated by 90°; and 2) a transversal raster pattern where layers were the print perpendicular to the length of the tensile specimen. In the case of the transversal specimen, there was no perimeter shell and it is noted that Torrado [16] noted this print raster pattern to be equivalent in terms of UTS with tensile test specimens printed in the ZXY print orientation.

Comparing the sensitivity to raster pattern between PLA and the PLA/NaCl composite revealed the addition of NaCl to reduce the difference in UTS between thetensile918transversal and crosshatched raster pattern (Fig. 6.). More notable is that this decrease in mechanical property anisotropy did not come at the expense of overall yield strength as has been the case in work related to decreasing the anisotropy of ABS-based ternary polymeric blends as explored in Torrado et al. [17].Fig. 5. a) non functionalized jute fiber and b) NaCl functionalized jute fiber. Note the presence of fibrils for the functionalized plant fibers.Biodegradabilitythe biodegradability of the PLA based composites evaluated in this study. Samples were buried for two and twelve weeks in topsoil which was moistened with food grade vinegar.

Samples were buried for two and twelve weeks in topsoil which was moistened with food grade vinegar. The effect of exposure to the soil environment was assessed by tensile testing of the exposed samples as well as the monitoring of weight change before and after exposure. No observable difference to weight was observed for any of the samples buried in the soil environment for two weeks. However, after 12 weeks, PLA specimens exhibited an increase in weight, while (1.95± 0.02 g increasing to 2.03 ± 0.01 g) the PLA/NaCl composite exhibited a decrease in weight (2.49 ± 0.03 g decreasing to 2.24 ± 0.12 g). In terms of tensile strength, both PLA and the PLA/jute fiber composite exhibited an increase in tensile strength compared to non- exposed control samples while the PLA/NaCl composite exhibited a decrease in tensile strength as compared to non-exposed samples (Fig.

7.) Further exploration is needed to understand if the decrease in tensile strength and tensile strength is due to biodegradation or the effect of moisture. NaCl is water soluble and the decrease in weight and tensile strength could be due to dissolution of this substance within the PLA matrix. The increase in tensile strength and weight for PLA asAnattempttoassessFig. 6 The addition of NaCl to PLA decreases mechanical property sensitivity to print raster orientation.919well as the increase in tensile strength for the PLA/jute plant fiber composite is also a finding that merits further study. Future work will include subjecting samples to biodegradability testing following ASTM D6954, a standard for assessing the biodegradability of polymers.Figure 7. The effect of moist soil exposure to PLA, PLA/jute fiber composite and PLA/NaCl composites.ConclusionsTailoring the physical properties of the biopolymer, PLA is possible through the addition of sustainable additives.

Here, when melt compounding was used to combine PLA with jute plant fiber, organo nanoclay pigments, and NaCl a discernable difference in physical properties was observed as compared to PLA alone. Explored in greatest detail was the effect of the addition of NaCl to PLA. Fractography of tensile test fracture surfaces revealed NaCl to adhere well to the PLA matrix as there was no evidence of particle pull out. The robustness of the bond between NaCl and PLA was also displayed by evidence of crack front propagation shared between the NaCl particle and the surrounding matrix. Further exploration of the benefit of combining PLA with NaCl was observed through the functionalization of jute plant fiber with aqueous NaCl. The use of palygorskite-based organonanoclay pigments also shows potential in playing a role in future PLA based 3D printable material systems.

The use of palygorskite-based organonanoclay pigments also shows potential in playing a role in future PLA based 3D printable material systems. Several aspects of this work bear the need for future study such as the increase in strength observed after exposing PLA and PLA/jute plant fiber composites to a moist soil environment for durations of two and twelve weeks.AcknowledgmentsThe work presented here was carried out in the W.M. Keck Center for 3D Innovation and the Department of Metallurgical and Materials Engineering, both at The University of Texas at El Paso (UTEP). The authors are grateful for funding provided through the AFOSR YIP program under grant number FA9550-14-1-0260 and the UTEP Campus Office of Undergraduate Research Initiatives (COURI).References[1] B. Stephens, P. Azimi, Z. El Orch, and T. Ramos, “Ultrafine particle emissions from desktop3D printers,” Atmospheric Environment, vol. 79, pp. 334–339, Nov. 2013.[2] K.-L. G. Ho, A. L. Pometto, A.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEvaluation of a Cyber-Physical Attack Effectiveness in Metal Additive Manufacturing by Selectively Modifying Build Layer ThicknessPatricio E. Carrion1, 2, Lynne M.

Unintentional misconfiguration of these parameters is easily detectable as it impacts the entire build. In this paper, however, we consider the case of a deliberate sabotage attack which causes misconfiguration localized to only few strategically selected layers. We propose a method on how such targeted misconfigurations can be executed without hacking into the firmware. Specifically, we altered a build file to mimic localized layer thickness modification by disabling laser beam exposure, while maintaining geometrical and visual part integrity. For two distinct laser powder bed fusion (L-PBF) systems and two metal alloys, we validated empirically the impact of such attack on part quality and demonstrated that it can avoid detection by non-destructive techniques (NDT).

The conducted attack illustrates susceptibility of AM to deliberate sabotage attacks and motivates the need of security solutions for this increasingly adopted manufacturing technology.Keywords: Additive Manufacturing; Laser Powder Bed Fusion; AM security, Cyber-Physical Attack; Sabotage.IntroductionAdditive Manufacturing (AM) has numerous advantages, such as ability to manufacture parts of complex geometry, reduced lead times and material waste, and on-demand manufacturing of spare parts at a location where they are needed [1–3]. Therefore, AM is increasingly adopted by various industries, including the aerospace, defense, and biomedical sectors.NIST classifies AM as a Direct Digital Manufacturing (DDM) technology, meaning that it fabricates physical objects from a digital design file using computer-controlled processes with little to no human intervention [4].

Because it relies so heavily on digital input files, in addition to the241digital control systems operating the AM machine itself, its security has been identified as a critical gap [4,5]. Security issues include technical data theft, sabotage, and illegal part manufacturing [6].This paper focuses specifically on cyber-physical sabotage attacks, i.e., attacks which originate as a manipulation in the cyber domain and cause effects in the physical domain [7]. It should be noted that the effects of interest to an adversary are rarely achieved immediately but are rather a product of a complex causal chain of effect propagations [8]. In the AM context, sabotage attacks can target properties of manufactured parts (and thus of systems into which they are integrated), the AM machine itself, or the manufacturing environment [9].

In the AM context, sabotage attacks can target properties of manufactured parts (and thus of systems into which they are integrated), the AM machine itself, or the manufacturing environment [9]. The dr0wned study [10] has demonstrated that sabotage of a part is rather a means and not the end goal per see; in the study, a sabotaged propeller broke after a brief operation time causing fall and destruction of the quadcopter UAV on which it was installed.Sabotage attacks are of especial importance for metal parts, because they are used in safety- critical systems. Currently, the most dominant AM process to manufacture net shape metal parts is Laser Powder Bed Fusion (L-PBF). Part sabotage on a L-PBF is the focus of this paper. We propose a novel method on how layer thickness, an essential PBF process parameter, can be modified locally without the need to compromise the 3D printer firmware.

We propose a novel method on how layer thickness, an essential PBF process parameter, can be modified locally without the need to compromise the 3D printer firmware. We employ this method to simulate an attack that selectively modifies a part’s layer thickness using two difference L-PBF systems with different powder feedstocks, 316L and 17-4PH Stainless Steel (SS). We then verify both the ability of this attack to remain undetected by an X-ray computer tomography (CT) and to degrade part’s tensile strength. Finally, we discuss the implications of the studied attack on the security posture in AM.Related WorkSeveral publications survey the state of AM Security field as a whole [6,11,12]. Analysis of software, firmware, and communication protocols used by desktop 3D printers has identified numerous vulnerabilities that can be exploited for their compromise [13].

Analysis of software, firmware, and communication protocols used by desktop 3D printers has identified numerous vulnerabilities that can be exploited for their compromise [13]. It has been demonstrated externally that external attackers can compromise manufacturing environment [10], run a malware on a compromised computer [14], hijack network communication with 3D printer [15], or compromise firmware of a 3D printer [16,17]. When compromised, any of these elements can then be used to conduct a broad variety of cyber- and cyber-physical attacks. Furthermore, several studies identified that the attack methods available to sabotage AM parts exceed those available for traditional manufacturing methods [18,19].Most AM sabotage attacks has been performed on Fused Deposition Modeling (FDM) [6], an AM process that is predominantly used with polymers and widely adopted in low-cost desktop 3D printers.

Introduction of random voids into a design file [14], selective substitution of print material through support structure material [20], manual modification of the design file [10], varying print orientation [20,21], changes of the extruded filament amount [17] or temperature [16], or even on-the-fly substitution of one print through an entirely different [17] – these are representative examples of methods used in part sabotage. All these attacks are of cyber-physical nature, i.e., manipulations in cyber domain case effect in physical domain[7]. The above- mentioned works focused on identification of new categories of attack methods that would lead to the degradation of part’s tensile strength or fatigue life.

The above- mentioned works focused on identification of new categories of attack methods that would lead to the degradation of part’s tensile strength or fatigue life. For composite material parts it has been shown that sabotage attacks can also be optimized to minimize deviations from the original design while achieving the performance degradation level set by an adversary [22].242Only few publications have investigated sabotage attacks of metal AM. Several theoretical analysis papers identified direct attack methods for sabotaging parts manufactured with L-PBF. These include manipulation of numerous process parameters [21], disturbance of network communication timing [23], and fluctuations of power supply to 3D printer [18]. In our related work [24], we investigated how manipulations of the Powder Delivery System (PDS) can be used to sabotage parts and verified experimentally that such attacks can degrade part’s fatigue life.

Manipulation of L-PBF scanning characteristics (laser power and scan speed) has been used to define a part’s failure point [25]. For the L-PBF systems integrating in-situ monitoring in a closed control loop to adjust laser power, it has been shown that manipulation of the sensor data can be used to impair the part’s quality indirectly [26].Experimental Set-upSelective layer thickness via build file manipulation Simulation and validation of the proposed method (i.e., localized layer thickness modification) was performed by fabricating and mechanically testing a set of round tensile specimens based ASTM standard E8 [27]. The geometry and dimensions of the tensile specimen is presented in Figure 1(a).

The geometry and dimensions of the tensile specimen is presented in Figure 1(a). The specimen can be destructively tested to experimentally evaluate the impact of the introduced layer thickness defect on function-related mechanical properties such as ultimate tensile strength and ductility (a detailed account of the tensile properties obtained is presented in the Test methodologies section).To introduce the layer thickness defect, we split the design of the specimen in three parts, top, bottom, and the cross-section located between the aforementioned parts (see Figure 1(b)). Since technically the design now consists of three individual parts stacked on top of each other, the process parameters for these parts can be selected/modified individually. We kept the default process parameters, including layer thickness, for both bottom and top parts.

We kept the default process parameters, including layer thickness, for both bottom and top parts. For the cross-section part, only the core laser exposure was set to zero power, meaning that the laser will still take time to traverse the scanning path while being turned off, therefore not melting powder. To conceal the attack and provide the appearance of a successful layer deposition, the contouring laser pass of the cross-section part was still performed with the recommended laser parameters. Furthermore, to maintain the overall specimen length (indicated in Figure 1(a)) while varying the cross-section part thickness, the bottom part size was kept constant while the top part was reduced in length.

As a result, the specimen’s critical location was targeted using the cross-section part, which may contain an arbitrary number of layers of unmelted powder.Build layout design & specimen fabrication To test the impact of the layer thickness defect on mechanical performance, we used three categories of specimen modification: (i) single part design, hereinafter referred to as “default”, (ii) a three part specimen design (discussed in the previous section and shown in Figure 1(b)) with a cross-section length corresponding to one layer thickness per the recommended process parameters of the powder feedstock being used, hereinafter referred to as “1-layer”, and (iii) a three part specimen design with a cross-section length corresponding to two times the layer thickness of the recommended process parameters, hereinafter referred to as “2-layer”. The former (default specimen) is used establish baseline mechanical properties.

The former (default specimen) is used establish baseline mechanical properties. The latter two (1-layer and 2-layer) are used to study impact of the introduced layer thickness defect.The build file created was used to fabricate specimens using two L-PBF systems, a dual- laser GE Concept Laser M2 (GE M2), and a single-laser EOS M290. Furthermore, two metal243powders were selected for each machine, 3l6L and 17-7PH stainless steel (SS), respectively. Both materials had similar powder size distributions ranging between 15–45μm and the same recommended layer thickness of 40μm. Extra details on other process parameters and powder feedstock properties maybe found in [24]. The specimens were manufactured vertically, using previously recycled metal powder. The recycling/reconditioning procedure used may be found in [28].

The specimens were manufactured vertically, using previously recycled metal powder. The recycling/reconditioning procedure used may be found in [28]. Schematics of the build layout, including the recoater arm direction and inter gas flow with respect to specimen location, for the GE M2 and EOS M290 systems are presented in Figure 2(a) and (b), respectively. Additionally, as can be seen in Figure 2(a), the specimens were divided in two groups, front (green colored) which were fabricated using one laser, while the ones located on the back (red colored), were fabricated with the secondary laser. For specimens fabricated with EOS M290, shown in Figure 2(a), the build layout was kept the same and it can be seen that the recoater direction and niter gas flow are perpendicular to each other.(a)(b)Figure 1. (a) Round tensile specimen size and geometry designed based on ASTM standard E8 [27].

(a) Round tensile specimen size and geometry designed based on ASTM standard E8 [27]. (b) Three- part split of the original design to simulate the selective layer thickness defect (i.e., layer skip location).After fabrication and built plate removal, the 316L SS specimens were tested without any post-processing, i.e., as-built surface condition. The 17-4 PH SS specimens were subjected to a CA-H1025 heat treatment procedure after we removed them from the built plate. This heat- treatment was selected to increase the strength of the material, which will provide a comparison between a ductile (316L SS) and brittle (17-4PH SS) material behavior [24].

This heat- treatment was selected to increase the strength of the material, which will provide a comparison between a ductile (316L SS) and brittle (17-4PH SS) material behavior [24]. No further processing244was performed after heat treatment; hence, 17-4 PH SS specimens were tested in the as-built surface condition.2-layerGas flowBack / Laser 21-layerDefaultRecoater direction2-layerFront / Laser 11-layerDefaultGas flow(a)w o l fw o l fs a Gs a G2-layerBack1-layerDefaultRecoater direction2-layerFront1-layerDefault(b)Figure 2. Schematic of build layout indicating specimen location with respect to the recoater arm and inert gas flow direction for the fabrications using (a) GE M2 with 316L SS and (b) EOS M290 with 17-4PH SS.

Schematic of build layout indicating specimen location with respect to the recoater arm and inert gas flow direction for the fabrications using (a) GE M2 with 316L SS and (b) EOS M290 with 17-4PH SS. The three specimen modifications are also indicated (default, 1-layer, and 2-layer).Test methodologies Monotonic tensile tests were performed according to ASTM standard E8 [27] and using an MTS Landmark servohydraulic system. All tests were conducted in displacement control mode at a rate of 0.012 mm/s; an extensometer with a 12 mm gage length was employed to record strain deformation up to 0.05 mm/mm. Two test per L-PBF system and specimen category (default, 1-245layer, 2-layer) were conducted. A tabulated summary including the feedstock material (316L/17- 4PH SS), specimen category, built plate location (front or back), and test results is presented in Table 1.

A tabulated summary including the feedstock material (316L/17- 4PH SS), specimen category, built plate location (front or back), and test results is presented in Table 1. The results include, the ultimate tensile strength, UTS, 0.2% offset yield strength, 0.2%YS, modulus of elasticity, E, percent area reduction, RA, and elongation to failure, εf.Results and DiscussionSpecimen comparison After fabrication, the efficacy of the cyber-attack was evaluated to see if it could go undetected by non-destructive inspection such as X-ray imaging. Comparison of X-ray reconstruction images of the default and 2-layer specimens at the area around the cross-section region, shown in Figure 3(a) – (d), demonstrate the different levels of severity of the cyber-attack.

It can be observed in Figure 3(a) and (c) that default specimens have internal defects that are characteristic to AM parts, but in the case of the 2-layer thickness modified 17-4PH SS specimen (see Figure 3(c)), the defects are unmistakably visible. Contrary to the aforementioned results, shown in Figure 3(b), the 2-layer thickness modified 316L SS specimen does have clear indications that the part has been compromised, but it was observed that defects were generally larger within the scanned area (in comparison to the “default” counterpart) and some were agglomerated near/at the modified layers.These results indicate that quality inspection of AM parts quality should be more rigorous and versatile since is seems that effectiveness and detectability (via non-destructive evaluations like X-ray tomography) of this type of attacks hinge on the AM system and material properties.

The reason being that for a given AM system, the melt pool depth prescribed by the manufacturer’s recommended processes parameters vary from material to material, leading to different laser power and layer thickness requirements [24,29]. Finally, it should be emphasized that the precise location of the modified layer(s) was known as well the specimen size being relatively small in comparison the actual parts. In realistic applications these conditions will most likely not hold, thus significantly reducing the probability of the attack detection.(a)(b)246ZZYYXX(c)(d)Figure 3. X-ray tomography images comparing the internal porosity of a (a) default and (b) 2-layer modified thickness 316L SS specimen, as well as (c) default and (d) 2-layer modified thickness 17-4PH SS specimen near/at the cross-section volume [24].Tensile behavior The stress-displacement response of L-PBF 316L SS default, 1-layer, 2-layer specimens are presented in Figure 4.

Since specimens were manufactured using a GE M2 dual laser AM system, specimens fabricated with laser 1 (i.e., front specimens) are denoted with the solid lines while specimens fabricated with laser 2 (i.e., back specimens) are denoted with dashed lines. The linear portion of the curves (which are related to the modulus of elasticity, E) do not vary from each other regardless of specimen category (i.e., default, 1-layer, 2-layer). This was further corroborated by comparing the modulus of elasticity, E, values from each test, presented in Table 1, where it was observed that values were similar (< 5 GPa difference). However, after yielding the specimens fabricated with laser 1 had higher yield stresses in comparison to those fabricated with laser 2, regardless of specimen category.

However, after yielding the specimens fabricated with laser 1 had higher yield stresses in comparison to those fabricated with laser 2, regardless of specimen category. Moreover, the effect of the attack was only evident after specimen failure, where it was observed that the area reduction, RA, and elongation to failure, εf, of 1-layer and 2-layer thickness specimens were considerably lower than the default specimens. Hence, the attack significantly affected the ductility of the part.

Hence, the attack significantly affected the ductility of the part. The reduced ductility of 1-layer and 2-layer thickness specimens was confirmed to be due to the cyber-attack rather than other variables because the failure occurred exclusively at cross-section region, while for the default specimens occurred elsewhere along the specimens reduced section.In the case of the default, 1-layer, 2-layer L-PBF 17-4PH SS specimens, presented in Figure 5(a) and (b), the influence of the attack was immediately evident as both specimen categories (1-layer and 2-layer) failed prematurely and before reaching yield. As can be seen in the close-up view of the stress-displacement curves shown in Figure 5(b), 1-layer and 2-layer specimens did not surpass a stress value of 600MPa, which is close to half the yield stress of the material.

Moreover, it is observed that the stress-displacement curves of the 1-layer and 2-layer specimens are not similar to those of the default specimens; thus, indicating that the stiffness of the parts were affected as well, i.e., the modulus of elasticity was lower (see Table 1). Finally, it is important to mention that these tensile properties of the default specimens were in agreement247with similar studies on the mechanical behavior of vertically fabricated L-PBF 3l6L and 17-4 PH SS parts [30,31].Table 1.

(a)Stress vs. displacement curves, as well as an (c) enhanced view plot comparing the tensile deformation behavior of default, 1-layer, and 2-layer L-PBF 17-4PH stainless steel specimens fabricated using a EOS M290 AM system.Discussion of security implicationsThe experimental results clearly verify our assumption that the proposed method to introduce the layer thickness defect can be used in cyber-physical sabotage attacks on metal AM parts. We also demonstrated that this category of attacks can be implemented by simple modification of the build design files, which significantly lowers the technical sophistication needed and consequently increases the potential exposure to such attacks. We have also shown that both the impact of such attacks and their detectability are material dependent.

We have also shown that both the impact of such attacks and their detectability are material dependent. From the249security perspective, it means that indiscriminately attacks might not necessarily yield significant impact, while targeted attacks would require participation of a subject matter expert and substantial knowledge about a part to be sabotaged – factors that reduce the security exposure. Lastly, at least on the designed specimens, this attack can be readily identified during quality inspection if the region of interest is known (see Figure 3). This means that even though material and time have been expended on the attacked build, the potential impact on the target system in which the part should be integrated can be easily averted.Conclusion and Future WorkAdditive Manufacturing (AM) is increasingly adopted, while its degree of computerization exposes it to a variety of cyber- and cyber-physical attacks.

The study of such attacks and their characteristics is a necessary prerequisite for the development of efficient countermeasures.In this paper, we focused on sabotage attacks against metal AM parts manufactured on a L-PBF machine. Specifically, we proposed a novel method to selectively modify process parameters at the layer thickness granularity. We demonstrated how this attack can be executed by simple modification of design file. We also evaluated the impact of the attack on a part’s function by conducting tensile destructive tests on specimens manufactured with two different SS alloys, 316L and 17-4 PH.

We also evaluated the impact of the attack on a part’s function by conducting tensile destructive tests on specimens manufactured with two different SS alloys, 316L and 17-4 PH. The experimental evaluation clearly shows that the attack is very effective in degrading the part’s performance by at least factor of 6 when comparing the ductile properties of 17-4PH SS.While we have expected that the defect size (or heights of the cross-section impacting the modified layer thickness) will have a direct proportionate impact on the failure characteristic, the results presented show that it is not necessarily the case. For example, for both 316L and 17-4PH materials, the 2-layer specimens were relatively more ductile than the 1-layer, which is counterintuitive as it was expected that thicker layers will lead to more defects.

The underlying mechanisms leading to these results will need further future investigation.The sabotage attacks modifying AM process parameters have, so far, relied on compromised firmware; however, the technical sophistication needed to compromising firmware is higher. The attack presented in this paper can be conducted via compromised computer or computer network. As hacking into computer systems and hijacking computer network communication can be conducted by readily available exploits and tools, the technical entry level of the considered attack is fairly low. To counter this, well-established cyber-security measures should be used to ensure the integrity of the design and build files.In our future work we plan to investigate how the proposed approach can be used to introduce different modification of the manufacturing process parameters at the layer granularity.

While our primary objective for this paper was to study a specific sabotage attack, a similar method can be used to introduce layer graded microstructural characteristics – a goal that might be of interest to benign applications as well.250AcknowledgmentsThis work was funded in part by the U.S. Department of the Navy, Office of Naval Research under Grant N00014-18-1-2488, in part by the U.S. Department of Commerce, National Institute of Standards and Technology under Grant NIST-70NANB19H170, and partially supported by the National Science Foundation (NSF) under grant #1919818251References [1] Shamsaei N, Yadollahi A, Bian L, Thompson SM. An overview of Direct Laser Deposition for additive manufacturing; Part II: Mechanical behavior, process parameter optimization and control. Addit Manuf 2015;8:12–35. https://doi.org/10.1016/j.addma.2015.07.002.[2] Yadollahi A, Shamsaei N.

aluation of a digitised splinting approach with multiple-material functionality using Additive Manufacturing technologiesA. M. Paterson, R. J. Bibb, R. I. CampbellLoughborough Design School, Loughborough University, Leicestershire, LE11 3TU.AbstractThe design and fabrication of custom-made wrist immobilisation splints can be a laborious process. In addition, patient adherence in terms of wear duration and frequency may be affected by a range of contributing factors including poor aesthetics, hygiene issues and fit. This paper suggests the use of additive manufacturing (AM) in a bid to resolve factors affecting adherence and improve the efficiency of design and manufacture. Particular attention is paid to the exploitation of multiple-material capabilities using Objet Connex technologies, with the intent to integrate completely novel and state-of-the-art characteristics within splints.

However, in order to exploit the many benefits of AM for customised splint fabrication, appropriate Three Dimensional (3D) Computer Aided Design (CAD) methodologies must be considered for splint design. Furthermore, a specialised CAD approach must be developed for splinting practitioners to allow them to create such geometries. As a result, this paper describes the development of a customised 3D CAD methodology for splinting practitioners to design custom-made splints, in order to evaluate such novel features only available through AM fabrication.IntroductionRheumatoid arthritis (RA) is a chronic, systemic autoimmune disease, which typically affects joints within the hands, wrists, ankles and feet (Biese, 2007; Melvin, 1982). Symptoms can include inflamed synovia and tendon sheaths and destruction of cartilage and bone, resulting in pain and discomfort (Melvin, 1982). Borenstein et al.

Symptoms can include inflamed synovia and tendon sheaths and destruction of cartilage and bone, resulting in pain and discomfort (Melvin, 1982). Borenstein et al. (1993) state that the approach to RA treatment is a multi-layered pyramid, consisting of ‘education, physical and occupational therapy, rest and nonsteroidal anti-inflammatory drugs” (pp.545). Occupational therapy in particular addresses limitations that patients may encounter during everyday activities in an attempt to circumvent the limitations, and to improve wellbeing and quality of life. One method of intervention is splint prescription; the perceived benefits of which are multi-faceted (Melvin, 1982; Taylor et al., 2003; Jacobs, 2003; Colditz, 1996): relieves pain through immobilisation and protection of affected joints. protects painful contractures from impacts, scarring and excessive movement. promotes movement of stiff joints through immobilisation of more mobile joints.

protects painful contractures from impacts, scarring and excessive movement. promotes movement of stiff joints through immobilisation of more mobile joints. encourages healing of fractures and contractures. prevents or corrects deformities and contractures. rests affected joints. provides support.i. ii. iii. iv. v. vi. vii.Practitioners such as occupational therapists and physiotherapists may prescribe either custom-made or ‘off-the-shelf’ prefabricated splints.

provides support.i. ii. iii. iv. v. vi. vii.Practitioners such as occupational therapists and physiotherapists may prescribe either custom-made or ‘off-the-shelf’ prefabricated splints. However, this paper will describe and discuss the design and fabrication methods of custom-made static wrist immobilisation (SWI) splints in particular (shown in Figure 1), since they are one of the most commonly prescribed splints amongst a range of conditions (Stern, 1991).656Flared edge to prevent forearm pinchingFolded edge around the thenar eminenceFolded edge at distal palmar creaseFigure 1: Traditionally manufactured static wrist immobilisation splint and common characteristicsTherapists must take into account the likelihood of their patients adhering to their splint wearing regimes. Unfortunately, Sandford et al. (2008) found that two-thirds of patients reported non-adherence, and advise therapists to be aware of and acknowledge low adherence levels.

Unfortunately, Sandford et al. (2008) found that two-thirds of patients reported non-adherence, and advise therapists to be aware of and acknowledge low adherence levels. There are several reasons for poor patient adherence in terms of wear duration and frequency (Melvin, 1982; Taylor et al., 2003; Callinan and Mathiowetz, 1996; Sandford et al., 2008; Spoorenberg et al., 1994):The splint does not address the patient’s condition.  The patient has received insufficient information about their condition.  The patient has not had sufficient information justifying the need for the splint.  The splint is unattractive.  The splint is difficult to don/doff.  The splint is uncomfortable to wear.  The splint is impractical in certain environments, or for certain tasks.

 The splint is unattractive.  The splint is difficult to don/doff.  The splint is uncomfortable to wear.  The splint is impractical in certain environments, or for certain tasks.  The patient may not be interested or informed on the potential beneficial outcome of wearing their splint.Hygiene issues may also contribute to low adherence; open-cell padding within splints can absorb moisture, such as perspiration, which can lead to odours and collection of bacteria (Coppard and Lynn, 2001). More generally and in terms of assistive devices, Louise-Bender Pape et al. (2002) reviewed literature linking perceived social stigma to the association of assistive devices, and can contribute to non-use of such devices. In addition, 78% of 27 participants reported immobilisation splints as unwieldy (Spoorenberg et al., 1994).The splinting processFigure 2 depicts the splint design and fabrication workflow, deduced from Lohman (2001) and Jacobs and Austin (2003).

The therapist must choose the Low Temperature Thermoplastic (LTT) based on a variety of properties prior to splint forming, including contour conformability, thickness and colour. Colour selection by the patient may be encouraged in a bid to improve adherence (Coppard and Lynn, 2001). However, choice is often dictated by stock available in clinics or properties that are deemed more important by the therapist, such as thickness. In addition, therapists are advised to consider and apply657padding before forming the LTT, to avoid inducing pressure to prone areas (Taylor et al., 2003; Coppard and Lynn, 2001; Cooper, 2007). Padding types can vary from sheets and stockinettes to silicone gel discs or pads, which can be placed over bony prominences for cushioning (Coppard and Lynn, 2001).

Padding types can vary from sheets and stockinettes to silicone gel discs or pads, which can be placed over bony prominences for cushioning (Coppard and Lynn, 2001). Other common SWI splint characteristics are proximal edge flaring to avoid forearm pinching, and folding of both the distal and thenar edges to provide a more comfortable edge against the skin whilst adding rigidity to the splint (Figure 1).Figure 2: Traditional splint design and fabrication workflow Deduced from Jacobs and Austin (2003) and Lohman (2001)658Although many of the characteristics described are integrated to address functional needs, results may appear unwieldy, voluminous and unsightly. For example, if cavities are integrated over bony prominences to relieve pressure, the appearance of the altered topology can be compromised. In addition, folded or rolled edges add volume to the palmar region, potentially affecting palmar grasp capacity.

In addition, folded or rolled edges add volume to the palmar region, potentially affecting palmar grasp capacity. In terms of the fabrication process, there is little room for error; the farther the therapist progresses down the fabrication workflow, the more challenging adjustments may be to make. If additional or replacement splints are required the whole process must be repeated.Additive Manufacturing for upper extremity splint fabricationSeveral research studies have explored the use of Additive Manufacturing (AM) for upper extremity splint fabrication. Fraunhofer IPA fabricated a prototype SWI splint using an EOS P100 Laser Sintering (LS) system with Polyamide (PA2200) powder (Breuninger, 2012) (Figure 3); the benefits were improved aesthetics, bespoke fit and integral fasteners to exploit the AM capabilities of part consolidation.

A hand immobilisation (HI) splint prototype developed by Materialise (fabricated using an EOS P730 LS machine with PA2200 powder) gave a reduction in weight (Pallari, 2011). Oxman (2011) fabricated ‘Carpal Skin’ using Objet Connex500TM technologies, by exploiting the system’s capabilities to enable multiple- material builds (Objet Ltd., 2012b). The structures within the splint-like gloves were dictated by a predefined pain map to allow or restrict movement via a reaction-diffusion pattern (Oxman, 2011). Not only does the use of AM in the context of upper extremity splinting allow for exploitation of bespoke fit and function, but both prototypes by Oxman (2011) and Fraunhofer IPA also demonstrated improved aesthetics as a result of geometric complexity and freedom; synonymous characteristics which are only viable as a combination through AM.Figure 3: Bespoke-fitting wrist splint fabricated using Selective Laser Sintering (EOS P100 system and Polyamide12/PA2200 powder).

Image Courtesy of Fraunhofer IPA, Stuttgart.659Sketches shown in Figure 4 by Bibb (2009) show the intent of AM for splint fabrication with lattice feature integration. The lattice structures were intended to look aesthetically pleasing whilst promoting airflow to the skin in a bid to reduce perspiration. An additional intent was to enable patients to personalise their splints by choosing their own perforation patterns in an attempt to improve patient adherence.Figure 4: AM splint conceptsThe investigators were also interested in exploring additional and novel approaches for multiple-material integration into splints, particularly the use of elastomeric regions for a variety of new design features as a result of Objet Connex capabilities. For example, radial and/or ulnar-based elastomer hinges could be integrated into the splint to allow for easier donning and doffing.

For example, radial and/or ulnar-based elastomer hinges could be integrated into the splint to allow for easier donning and doffing. In addition, specific elastomer regions could be placed over bony prominences (similar to gel discs and pads in traditional splinting). Both concepts could be achieved without affecting the overall topology of the splint. It was anticipated that regional patches could protect and cushion bony prominences, whilst providing dynamic pressure over areas prone to fluctuating oedema with an aim to prevent oedema pooling.However, if such a design and manufacturing approach is to be realised for upper extremity splinting in a clinical domain, supporting software technologies need to be developed and tailored to suit the target user. Pallari et al. (2010), Rogers et al.

Pallari et al. (2010), Rogers et al. (2000) and Knoppers and Hague (2006) support the need for product-specific Computer Aided Design (CAD) software for practitioners, to enable the intended users to adopt quickly and easily without dedicating time to learn. In addition, Smith (1991) justifies the need for customised software, as it gives the intended users the tools to achieve what they want and need whilst removing unnecessary functions.Research Aim and objectivesA number of weaknesses in traditional splinting have been identified, as well as the strengths and feasibility of AM in splint fabrication. The need for customised Three- Dimensional (3D) CAD software to support the use of AM has also been noted. Therefore,660the aim of this paper is to evaluate the feasibility of new design features for upper extremity splinting only made viable through AM, as well as the digitisation of traditional splint characteristics. Objectives to achieve the aim are: Objective 1.

Objectives to achieve the aim are: Objective 1. The replication of key splint features/characteristics in a virtual environment. Objective 2. Exploration of new features only viable through AM, which could potentially address concerns regarding aesthetics, hygiene, comfort and form. Features include; 2.1. Aesthetically appealing and personalised lattice integration. 2.2. Multiple-material integration (using Objet Connex technologies) to;2.2.1. Imitate gel-discs/pads. 2.2.2. Potentially replace the need to create bulbous features in splints. 2.2.3. Provide dynamic pressure over areas susceptible to fluctuatingoedema.2.2.4. Integration of flexible hinges for easier donning/doffing.Objective 3. Evaluation into the feasibility of the digitised splinting approach via aspecialised CAD software prototype; 3.1. Refine replicated characteristics into a virtual workflow. 3.2. Devise a representation of the digitised approach via a softwareprototype.3.3.

Refine replicated characteristics into a virtual workflow. 3.2. Devise a representation of the digitised approach via a softwareprototype.3.3. Evaluate the digitised approach, specifically new features as a result ofAM.It should be noted that the scope of this investigation is purely related to the data manipulation of acquired patient scan data within 3D CAD in order to support AM splint fabrication. The investigation does not address data acquisition methods suitable for clinical environments, although Paterson et al. (2010) review data acquisition methods to generate test data in support of this investigation. Similarly, the process does not investigate the use of Finite Element Analysis (FEA), although it can be assumed that this would be a crucial feature if the digitised approach were to be realised.

Similarly, the process does not investigate the use of Finite Element Analysis (FEA), although it can be assumed that this would be a crucial feature if the digitised approach were to be realised. It should also be noted that the digitised splinting approach is not intended to make the splinting profession obsolete, nor does it imply the redundancy of splinting practitioners. The intended proposition is to provide therapists with an additional toolset to capture their design intent quickly and easily, whilst addressing the needs and concerns of patients through improved aesthetics, fit and functionality.A systematic approach was adopted to develop a sequence of functions to meet theobjectives.

Each objective will be addressed individually.MethodsObjective 1 – replication of key splint features/characteristics in a virtual environmentThe critical splint characteristics required for the specialised splinting workflow were tested iteratively in a range of 3D CAD software packages, including Geomagic Studio (2010) and McNeel Rhinoceros (2011), using an action research strategy. The approach was intended to test and refine different CAD tools and CAD strategies to capture and replicate the design intent of typical splint features, such as pressure relief cavities. An example of the iterative testing is described by Paterson et al. (2011).661Objective 2 - Explore new features only viable through AM Objective 2.1 - Aesthetically appealing and personalised lattice integrationSimilar to addressing Objective 1, an action research strategy was adopted to explore lattice applications to a previously defined splint surface.

After iterative testing, it became apparent that an additional feature within the workflow was required to form a splint border (Figure 7, flow diagram item 3.1), which would encase the lattice structure. This would address inherent issues of sharp protrusions, which could otherwise cause lacerations if borders were not implemented into the proposed splints.Objective 2.2 - Multiple-material integration/heterogeneous structuresAs mentioned previously, the purpose of using Objet Connex technologies was to explore the viability of new features in the context of splinting, making the transition from homogeneous to heterogeneous splints. Objet Connex technologies can deposit materials in pre-defined combinations, which dictate regional material proportions to form so called Digital MaterialsTM (DM).

Objet Connex technologies can deposit materials in pre-defined combinations, which dictate regional material proportions to form so called Digital MaterialsTM (DM). For example, in order to create variations in shore hardness amongst the Objet flexible (FLX) FLX97-DM range, the primary material (TangoPlus) and secondary material (VeroWhitePlus) are interspersed by varying dual-jet distribution using Objet’s PolyJet Matrix™ Technology. The material depositions are then cured using ultra- violet (UV) light (Objet, 2012; Objet Ltd., 2012c).There were two main approaches to consider for multiple-material integration; Continuous Functional Grading (CFG) (Figure 5a) versus Stepped Functional Grading (SFG) (Figure 5b). CFG proposes a gradual transition of one or more properties across a defined volume, and has been of significant interest across a range of disciplines.

CFG proposes a gradual transition of one or more properties across a defined volume, and has been of significant interest across a range of disciplines. However, the implementation of such a feature in terms of CAD modelling strategies, file export and AM systems is still under development (Erasenthiran and Beal, 2006). Knoppers and Hague (2006) and Siu and Tan (2002) highlight the difficulties in representing functionally graded geometries in Boundary Representation (B-rep) modelling, due to the geometries being defined by a series of surfaces to determine the topology of a closed volume. If functional grading were to be a viable feature in splints, an alternative or additional modelling strategy within the CAD methodology (e.g. voxel-based modelling) would need to be adopted. A consequential limitation, however, is the increased memory power, processing consumption and modelling complexity during the design phase (Oxman, 2011; Erasenthiran and Beal, 2006).

A consequential limitation, however, is the increased memory power, processing consumption and modelling complexity during the design phase (Oxman, 2011; Erasenthiran and Beal, 2006). This would add unnecessary complexity to the geometry and CAD strategies. In addition, current Objet Connex technologies only offer a finite number of pre-defined deposition variations, which limit the CFG approach. As a result, CFG was considered excessively complex and unnecessary for this particular application, and it was decided that SFG would be sufficient to deliver the finite variation that Objet Connex technologies offer.The disadvantage of SFG is that it creates boundary lines between adjacent materials properties, which, depending on the differences in the adjacent properties could lead to stress concentrations at the boundary or abrupt changes in flexural modulus that could cause unwanted folds or creases at the boundary.

The worst-case scenario can be imagined to be a choice of only two materials, for example very rigid and very flexible. In that circumstance, the abrupt change at the boundary is likely to present problems. This would be mitigated by the ability to specify a greater number of regions and a sufficiently high number of materials property choices. It can be assumed that, for a given application, a sufficiently high number of materials choices could satisfactorily approximate continuous variation.662a.Continuous FunctionalGrading (CFG)b. Stepped Functional Grading (SFG)TangoPlus VeroWhite DM 9740 TangoPlus VeroWhite DM 9750 TangoPlus VeroWhite DM 9760 TangoPlus VeroWhite DM 9770 TangoPlus VeroWhite DM 9785 TangoPlus VeroWhite DM 9795Figure 5: Functional Grading versus Specific Boundary RepresentationObjective 2.2.1 - Imitate gel-discs/padsThis objective was to explore the possibility of using multi-material capability to reproduce the function of gel pads.

In existing practice, the rigid shell of the splint is raised in local areas to provide space to insert soft pads. These are required to reduce pressure on bony prominences and improve comfort. The ability to define material properties within a localised region enables the body of the splint to be made soft in the required area. This eliminates the pads and reduces bulk of the splint.Objective 2.2.2 - Potentially replace the need to create bulbous features in splintThis is essentially the same function as 2.2.1 except that in existing practice some areas of the splint may be raised to provide space for sensitive areas or bony prominences but without placing a gel disc/pad into the space.Objective 2.2.3 - Provide dynamic pressure over areas susceptible to fluctuating oedema This function is currently very difficult to accommodate with a single material rigid splint.

The ability to alter the flexibility of the splint, possibly in specific regions would enable a conformal fit whilst allowing some degree of expansion that would accommodate temporary or fluctuating swelling. This would be a unique advantage of the AM approach over existing practice.Objective 2.2.4 - Integration of flexible hinges for easier donning/doffingWith existing practice, the rigid splint has to incorporate large openings to allow donning and doffing. This leaves large regions unsupported. In addition, straps and fasteners have to be added into or onto the splint, which adds construction time, increases bulk and adds features that will be more susceptible to damage and wear. The ability to build in flexible joints or hinges enables a splint design with greater coverage and therefore increased support. The ability to build in hinges and fasteners reduces the bulk and the time required to add those features after forming the splint.

The ability to build in hinges and fasteners reduces the bulk and the time required to add those features after forming the splint. This approach has many other potential advantages in the elimination of purchasing and stocking fasteners, straps, etc. In addition, it is frequently straps and fasteners such as Velcro that wear, fail or become unacceptably soiled in use earlier than the splint itself.In terms of 3D CAD strategy testing to support the digitised splinting approach, the refined strategy for enabling multiple-material builds was to create a curve on the splint surface to act as a trim boundary prior to pattern application. After applying the pattern to the surface, the splint surface could then be trimmed (Figure 6a). Separate patches could then be thickened (Figure 6b,c), and exported as separate shells in a single STL file. Each shell can subsequently be assigned a digital material in the Objet preparation software.

Each shell can subsequently be assigned a digital material in the Objet preparation software. For example, in663Figure 6c there are three shells. The predominant grey shell could be a rigid material; the green shell could be an intermediate flexible material whilst the red shell could be very soft.a.b.c.Figure 6: Example of SBR integration over a bony prominence (red- pisiform) and medial/ulnar hinge (green) a. Three separate surfaces present b. thickened splint c. coloured regions of splints, depicting different materialsObjective 3 - Evaluation of approach via a specialised splinting CAD software prototypeThe requirements established after CAD testing were refined into a specialised CAD software workflow (objective 3.1) (Figure 7). The order was determined by consideration of the traditional splinting workflow depicted in Figure 7, as well as constraining/best practice CAD approaches (e.g. detailed application to surfacing prior to thickening).

detailed application to surfacing prior to thickening). An important feature to note in the digitised approach was the intent for traversal operation; the ability to move back and forth between different features independent of one another. This, fundamentally, was anticipated as an important step for splint fabrication when compared to traditional splint fabrication, where the further one progresses through design and fabrication, the more difficult and time consuming it may be to make adjustments.664Figure 7: Refined digitisation workflow * = Existing feature/activity ** = New feature/activity ***= Necessary feature as a result of a new feature/activityHaving established a refined workflow, a high-fidelity concept software prototype was developed within Microsoft® Access® (2010) using Microsoft® Visual Basic® for Applications (2010), which could be used to depict the workflow for final evaluation (objective 3.2).

The prototype featured a viewport, and a series of slider controls and drop- down boxes to control features such as pattern density (Figure 8). When controls were adjusted, the image in the viewport would change to suit the recent adaptation to provide direct visual feedback to the user. The prototype also featured a pattern library and a separate fastener library, which allowed users to browse different perforation patterns and fasteners. The user would then be able to select the desired options from the libraries for automatic placement to the splint geometry in the viewport.665Figure 8: Prototype interface example, and resulting changes from altering pattern perforation densityTen splinting practitioners within the UK were invited to participate in evaluation sessions (physiotherapist: n=2, occupational therapist: n=8). Eight participants had one-to-one sessions, whilst two participants joined in one session for their convenience.

Eight participants had one-to-one sessions, whilst two participants joined in one session for their convenience. Each evaluation session comprised four activities; a briefing into the intent of AM for splint fabrication; a666demonstration of the prototype; user trials of the software prototype and a semi-structured interview. Qualitative data were captured regarding the digitised approach, together with specialist feedback regarding multiple-material into splints. Participants completed a demographic questionnaire to gather data including the number of years splinting experience. In addition, participants signed informed consent forms agreeing to the capture of audio and computer screen-capture recordings whilst testing the prototype. The recordings were used to identify trends in opinions and were transcribed into NVivo (2011), for coding and trend identification.

The recordings were used to identify trends in opinions and were transcribed into NVivo (2011), for coding and trend identification. In addition, proof-of-concept splints were manufactured using various AM processes to give physical, tactile representations of the intended output (Figure 9). The purpose was to show participants what could be achieved using AM, to ensure that the intent was conveyed effectively to participants in the context of upper extremity splinting.integrationUlnar hingeCushioning regiona.b.c.Figure 9: Proof-of-concept AM splintsa. Fabricated using 3D Systems SLA® 250 (3D-Systems Xtreme material).b. Fabricated using Objet c. Fabricated using Objet Connex500 TM, with multi-material regions (ulnarConnex500TM (RGD5160-DM).hinge: FLX9760-DM; cushioning region: FLX9740-DM; remainder of splint: VeroWhitePlusTM FullCure 835).Results and discussionA wide range of positive comments and suggestions were gathered during the evaluation sessions.

The approach not only highlighted many exciting new avenues for exploration within splint design and fabrication, but it also expanded the window of opportunity for Objet Connex technologies. Participants mentioned that the integration of multiple materials within a single splint was a completely new and exciting toolbox for therapists, which could not be achieved using existing fabrication methods. Participants were interested in new applications as a result of multiple-material capabilities, such as protection and cushioning of bony667prominences without compromising the splint topography. Of the three participants who were asked, all three participants favoured ‘mono-splints’ with elastomer hinges (as shown in Figure 9c), to make donning and doffing easier for patients with restricted dexterity; two participants even suggested having two hinges, placed along the radial and ulnar aspects of the wrists to assist further.

Three participants were interested in using elastomer materials to apply dynamic pressure over areas prone to fluctuating oedema. In extension to these applications, one participant suggested the use of elastomer materials in splints for treatment/management of burns and scars, and protection of post-surgical metalwork (i.e. alignment screws and brackets). Another participant suggested integration of elastomer borders on splint edges, which could aid in pressure distribution and provide a softer interface between the more rigid structure of the splint and the skin. Therapists also expressed their interests in allowing partial dynamic features to allow or restrict movement, similar to Carpal Skin by Oxman (2011).Many participants suggested additional creative features to be integrated within the digitised workflow, such as multicolour integration; one participant asked whether photographs could be applied to splints.

At present, colours are limited on the Objet Connex; 3D Systems offer the ZPrinter® 450 multicolour 3D printer (3D Systems Inc., 2012), but this would result in compromising multiple-material feature integration. However, Holmes (2012) speculates that Objet Connex Technologies will soon offer multicolour functionality in addition to multiple-material printing; Oxman (2012) prompted this speculation as a result of recent work, but the intent for market release is yet to be confirmed. Therapists were also keen on part consolidation with regards to fastener integration. Participants felt that the workflow depicted in the software prototype demonstrated the workflow used in traditional fabrication, with the exception of new features, and placement of colour and thickness controls.

In response to traversal movement within the workflow; eight participants felt this was a useful feature, particularly since variables could be altered independently from one another; a desirable novelty when compared to traditional splinting.The intent of edge filleting in the CAD workflow (Figure 7, flow chart item 4.4) was to relieve the need for rolling of splint edges, consequently reducing cumbersome structures presented in traditional splinting. Therapists were surprised at the strength and rigidity of the prototype splints without having folded or rolled edges within the palm. One participant was interested to know whether this would improve palmar grip capacity, since material volume was reduced.

One participant was interested to know whether this would improve palmar grip capacity, since material volume was reduced. However, some participants felt they would still want the opportunity to produce the effect of rolled and folded edges within the splinting software.Despite the abundance of positive responses, participants also raised concerns regarding data acquisition methods, structural integrity, fabrication time and cost of the approach. Two participants were concerned that oedematous, inflamed regions might protrude through the lattice structure, potentially causing discomfort for the patient. Another participant felt that patient adherence would not only be improved as a result of improved aesthetics, but the level of improved wear duration could potentially be detrimental; they were concerned that patients would be wearing their splints for longer durations than is necessary or recommended, potentially resulting in splint dependency.

Cooper (2007) states that static splints should not be worn more than necessary as they contribute to stiffness, atrophy and joint disuse. Therefore, a more stringent wearing regime might be required to combat this concern if AM splints were to be eventually prescribed.668Future workAlthough the investigation highlighted many new positive and exciting applications within AM splint fabrication, there is much to be done before the approach can be used as a clinically viable treatment method. To date, the software prototype described in this paper is purely a visual representation for evaluative purposes. Future development is required to create fully functional specialised splinting CAD software. New features must also be investigated for feasibility into workflow integration, such as edge-folding.

Future development is required to create fully functional specialised splinting CAD software. New features must also be investigated for feasibility into workflow integration, such as edge-folding. The ability to re- use previously defined splint files to make replacement splints must also be an integral feature in the CAD approach, as this would allow therapists to fabricate duplicate splints without incurring additional design time. The opportunity to adjust previously defined splint files should also be explored, so therapists may alter one or many design variables where required.

The opportunity to adjust previously defined splint files should also be explored, so therapists may alter one or many design variables where required. Alternative file format exports are also required to support depiction of complex organic geometry, whilst allowing for multiple-material and mono/multicolour capabilities; the development of STL 2.0 or Additive Manufacturing File Format (AMF) is promising, particularly for the composition of complex geometries and multiple-material builds in support of the Objet Connex capabilities (Hiller and Lipson, 2009). In terms of CAD process efficiency, Objet have developed CADMatrixTM; a plugin for material assignment within selected CAD applications prior to automated STL creation (Objet Ltd., 2012a). If the proposed software prototype were to be developed, this approach could be applied within the custom software.Cost was a concern expressed by the participants that demands further work.

If the proposed software prototype were to be developed, this approach could be applied within the custom software.Cost was a concern expressed by the participants that demands further work. It is too early in the development of the CAD approach to predict costs realistically. However, the advantages of the CAD/AM approach for future cost reduction is based on two premises. Firstly, there is very little scope for cost reduction in existing practice. The materials costs incurred in current practice are minimal and by far the greater proportion of cost is attributed to time and salary costs for the professionals involved. There is little scope for cost reduction of the existing materials and significant cost reduction would only be possible by speeding up work or reducing salaries, neither of which is likely to be acceptable and any savings would be small and incremental.

A shift to a CAD/AM approach would achieve cost savings by eliminating physical work (especially for remakes or replacements) and enabling a much faster design stage to be completed by the professionals. The separation of the manufacture from the design enables designs to be done at any time, potentially without the patient being physically present, whilst manufacture does not incur salary costs. Whilst AM machine and materials costs are currently high, it is reasonable to assume that these costs will reduce significantly over time as has already been demonstrated by other technologies. The potential advantages of a CAD/AM approach in the provision of custom-fitting medical devices have been recognised in other research, such as the provision of maxillofacial prosthetics (Bibb et al, 2010).Further development of medical grade materials conforming to ISO 10993 category standards will also be necessary for this application, in preparation for clinical trials.

In response to burns/scar treatment, further research into suitable materials would be required. The multiple-material splint shown in Figure 9c also suffered a significant amount of warping and deformation at room temperature; these issues will also need addressing.Although not within the focus of this investigation, FEA should also be investigated to address concerns regarding structural integrity; both the way in which it may be applied to the splint geometry in 3D CAD, but also the way in which the user would interact with the669feature. The approach would need to be easy to use with the potential for automated alterations through analysis and regeneration of problem areas. Therapists expressed the interest into multiple-material splints to either permit or restrict movement, termed ‘Synthetic Anisotropy’ by Oxman (2011).

Therapists expressed the interest into multiple-material splints to either permit or restrict movement, termed ‘Synthetic Anisotropy’ by Oxman (2011). Therefore, not only does research need to be performed into how to achieve such results in terms of 3D CAD modelling and specification of Objet Connex materials, but also the most efficient way to integrate such features within the specialised splinting CAD software.In response to therapists’ concerns into oedematous regions protruding through lattice structures, displacement mapping could be used on the outer splint surface to overcome such problems. This could be used instead of or in addition to material-to-perforation ratio alterations.

This could be used instead of or in addition to material-to-perforation ratio alterations. Not only would this aim to resolve issues related to oedema, but it would also be a step forward in resolving concerns regarding structural integrity.AcknowledgementsOur sincerest thanks to Lucia Ramsey at the University of Ulster, Belfast, Ella Donnison at Patterson Medical Ltd., Sutton-in-Ashfield, Jason Watson at Queens Medical Centre, Nottingham, and Dr. Dominic Eggbeer at The National Centre for Product Design and Development Research, Cardiff for their help, support and advice throughout the project. Extended thanks to all participants involved in the study, for their time, expertise and valued feedback. This research was funded by Loughborough University.3D at: ZPrinter® http://www.zcorp.com/en/Products/3D-Printers/ZPrinter-450/spage.aspx [Accessed June 18, 2012]. BIBB, R.J., 2009. Lattice within splints. Drawing. July 22. Loughborough University, Loughborough.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEvaluation of a feed-forward laser control approach for improving consistency in selective laser sinteringT. Phillips*, S. Fish*, J. Beaman*Department of Mechanical Engineering, University of Texas at Austin, TX 78712AbstractSelective Laser Sintering (SLS) is a popular industrial additive manufacturing technique for creating functional polymer components. One of the biggest limitations today with SLS is its poor mechanical consistency when compared with traditional manufacturing techniques, inhibiting the use of SLS among structurally critical components. Evaluation of the SLS process has revealed that the quality of components is strongly affected by the thermal history during the build process and poor control over this can lead to premature part failure.

This paper will discuss a novel technique of improving in-situ thermal control by implementing a feed-forward laser controller that uses dynamic surrogate modelling to predict optimal laser power to achieve desired thermal characteristics. Thermal and destructive testing results will be presented showing that the described laser power controller is capable of decreasing the standard deviations of post sintering temperature by up to 57% and ultimate flexural strength by up to 45%.IntroductionSelective Laser Sintering (SLS) is a subset of Powder Bed Fusion (PBF) additive manufacturing (AM) that uses a high-powered laser to fuse polymer powder into arbitrary 3D geometries whose features are defined in a 3D CAD file. This is accomplished by depositing a thin layer of polymer powder onto a build surface and scanning a laser across the surface to fuse the material in a shape resembling a 2D cross-section of the final 3D part.

Another thin layer of material is spread on top of the previous and the laser now scans the surface according to the next geometric cross-section, fusing this material to the previous layer. Layer-by-layer, the laser fuses the material until the component is completed, after which the un-sintered powder is removed and the 3D component is revealed. While many other AM techniques require supporting structures to ensure the components retain their shape, SLS is capable of creating free-floating components without the need for supports. This enables a user to pack a large amount of components into a single build.As with most AM techniques, SLS suffers from poor repeatability when compared with traditional manufacturing techniques. Small variations in strength and dimensions exist among additively manufactured components that may limit their usage.

Small variations in strength and dimensions exist among additively manufactured components that may limit their usage. Improving component consistency has been the subject of much research and it has been found that laser energy and processing temperature have a large influence on the final component strength [1] [2] [3] [4] [5] [6]. Many of the SLS machines in use unintentionally create thermal gradients on components, leading to the high variability of strength. These thermal gradients can be reduced by implementing a more precise heater system, but eliminating temperature differences across the1673powder surface is not possible with these techniques. An alternative approach is to modulate laser energy to account for whatever temperature differences exist on the pre-heated SLS powder. This paper will present a feed-forward approach to modulating laser power in order to reduce post- sintering temperature differences in components.

This paper will present a feed-forward approach to modulating laser power in order to reduce post- sintering temperature differences in components. Thermal information is collected in-situ and used in an adaptive surrogate model to predict optimal laser power, which is then sent to the custom laser controller. Testing shows that this technique greatly improves the consistency of components.SLS TestbedThe experiments presented in this paper were performed on ALM PA650 nylon material using the custom research SLS machine developed at the University of Texas at Austin known as the Laser Additive Manufacturing Pilot System (LAMPS). This system was developed for process control research and exposes all control variables to the user. The laser control technique discussed here makes use of one of LAMPS’ thermal cameras, a FLIR A6701. This Mid-Wave Infrared (MWIR) camera is located approximately 1 meter from the build surface at an angle of 28º.

This Mid-Wave Infrared (MWIR) camera is located approximately 1 meter from the build surface at an angle of 28º. It captures 640x512 pixels images at 60 Hz in the 150-300ºC temperature range and has a spatial resolution of approximately 400 μm per pixel. A blackbody is used in-situ for validating temperature values during runtime.A commercial galvanometer controller is used to control the laser power and position during normal operation of LAMPS. This controller, and every commercial galvanometer controller available at the time of writing, does not allow for intra-vector power changes, meaning obtaining the high level of control over laser power required for overcoming pre-heating thermal gradients is difficult. To overcome this issue, a custom solution was developed that works in conjunction with the commercial controller on LAMPS. This so-called Position-Based Laser Power (PBLP) controller sits between the galvanometer controller and the laser, as seen in Figure 1.

This so-called Position-Based Laser Power (PBLP) controller sits between the galvanometer controller and the laser, as seen in Figure 1. The PBLP controller receives the original laser power signal from the galvanometer controller as well as position feedback from the galvanometers and a description of the desired laser power surface from the LAMPS host computer. It uses the original laser power signal from the galvanometer controller as a trigger. When the signal is high, the PBLP controller continuously samples the galvanometer position, solves the laser power surface for that location, and outputs the correct signal to the laser. This allows the PBLP controller to have no knowledge of the laser scan path, yet still actuate the laser at the correct time intervals.1674LAMPSMark& Jump CommandsEClOOO ControllerXY2-100______ command.

_____ _Signa lD/A ConversionGa lva ControllerGalvaPositionPBLP ,-------------1 ~ - - - ' - - - -I:Analog Input Module:[ : ]Laser Power 1 ~ - - - - - - 1PBLP PWMLaser.,_ _____________ P_o_w_e_r_S_u_rf_a_c_e..,1,.11HS Digital 1/0 ModuleI :compactRIO1,_ - - - - - - - - - - - - ..Figure 1: Position-Based Laser Power ControllerLaser ControlA feed-forward laser control technique was developed in order to control the post-sintering temperature of components. Previously, researchers have developed in-situ monitoring techniques and real-time feedback systems for metal additive manufacturing [7] [8] [9] [10]. These systems are successful in homogenizing post-sintering temperatures of metal components, but would be challenging to implement in a polymer SLS system.

These systems are successful in homogenizing post-sintering temperatures of metal components, but would be challenging to implement in a polymer SLS system. The difficulty lies in the amount of thermal energy emitted at the processing temperature of polymers, as the amount of energy emitted at the melting temperature of nylon is two orders of magnitude lower than that emitted at the melting temperature of steel. This makes it difficult to find a sensor capable of the 10-20 kHz rate required for real-time control of the laser system [11]. Instead, a feed-forward approach is taken in this paper that allows for the high level of control desired without requiring custom sensors.In order to properly implement a feed-forward controller for laser power, a model is needed to relate laser energy deposited to temperature increase experienced by the powder. Since the SLS process is highly dynamic, the model should be able to adapt to changes that occur throughout the building process.

Since the SLS process is highly dynamic, the model should be able to adapt to changes that occur throughout the building process. For this, a dynamic surrogate model was used that relies on in-situ thermal data to model the SLS system. Due to differences in laser power transmission and powder conduction paths, it cannot be assumed that the same laser power produces the same effect on every area of the powder bed. Therefore, the powder bed is discretized into 1mm x 1mm areas, each with their own thermal model.The MWIR camera records images during each layer and the resulting information can be condensed into two useful images: pre-sintering and post-sintering temperature. The pre-sintering temperature image is simply the MWIR image taken before any lasing has occurred. The post- sintering temperature image is a composite created by taking the maximum value each pixel recorded during lasing.

The post- sintering temperature image is a composite created by taking the maximum value each pixel recorded during lasing. Subtracting the pre-sintering temperature from the post-sintering temperature gives the temperature increase experienced by the powder due to lasing. A perspective transformation and image registration process is performed on each of these images to transform them into the CAD coordinate system. This allows the laser scan files to be used as a mask to extract temperature information from each location that received laser energy, as seen in Figure 2. For each of the 1mm x 1mm areas on the powder bed that received laser energy, the corresponding1675thermal model is updated with the new laser power and temperature increase measurement. A weighted linear regression is performed on each model to determine the relationship between delivered laser energy and expected temperature increase for that area.

A weighted linear regression is performed on each model to determine the relationship between delivered laser energy and expected temperature increase for that area. Each new data point is added to its model with a high weight and each time a new measurement is added, all previous measurements have their weight decreased. This produces a forgetting factor that allows the thermal models to deviate from old information that may no longer accurately represent the system [12]. For the experiments discussed here, the forgetting factor was fixed at 0.8, providing a good balance between model stability and system tracking.

For the experiments discussed here, the forgetting factor was fixed at 0.8, providing a good balance between model stability and system tracking. This process of analyzing the thermal images and updating the thermal models takes place during the powder recoating process after each layer and adds a negligible amount of time to the build.a)E • ■ • Ill D :I ■ • ■ ■ ■ • • • ■ ■ ■ ■ ■ ■■ ■ ■ ■ ■ ■■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ li1 ■■ ■ ■ ■ ■ E 50mm 50 mm■ ■ ■ ■ ■ E 50mm 50 mmb)■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■ ■c)lie I... ..:r:lillii1Iiiii1IIIll • Iii! ~ Illp!!ii..

..:r:lillii1Iiiii1IIIll • Iii! ~ Illp!!ii.. , ■ • i..11 ~ i!! ■ • " t': ■ • 1!11 ■ ., l~•l~II:It\'!I,"I! ■t"" ~1.llaIIIIIllI! 11Figure 2: Images depicting (a) a composite post-sintering image after coordinate transformation into machine space, (b) thebinary mask created from scan file, and (c) the composite post-sintering image with binary mask applied.Prior to sintering each new layer, the pre-sintering temperature is measured and the thermal models are consulted to predict the optimal laser power. The laser scan file for the upcoming layer is used to determine which areas are to be lased and, consequently, which thermal models must be queried. For each area that is to be sintered, that area’s pre-sintering temperature and desired post- sintering temperature are given to the corresponding thermal model, which predicts the optimal laser power in that area.

This is repeated for each of the areas that the laser scan file indicated will be sintered on the upcoming layer. Once the optimal laser power has been predicted for each of these areas, a least-squares regression is performed to fit a parametric surface to the optimal powers and their machine coordinates. An example of such a surface is seen in Figure 3, where the red data are the predicted optimal powers and the blue data is the surface fit to them. The coefficients of this surface are sent to the PBLP controller, which modulates the laser power in real-time to match the desired power surface.

The coefficients of this surface are sent to the PBLP controller, which modulates the laser power in real-time to match the desired power surface. A block diagram of this control process can be found in Figure 4.1676Power Surface80~ a, 60 ~ II.401001001 001 00X(mm)Figure 3: Example power surfaceDesired TempSintering ModelBuild LayerPre-Sintering Thermal MeasurementPost-S intering Thermal MeasurementNew La erFigure 4: Automated laser control system block diagramExperimentsThe feed-forward laser control technique described above was evaluated by comparing the results it produced with baseline tests that used a pre-determined, fixed laser power regardless of pre-sintering temperature. The first test consisted of 99 flexure bars, 48 of which were built using the automated laser control (ALC) technique. These were built in groups of 10, with ALC and baseline specimens dispersed throughout each build.

These were built in groups of 10, with ALC and baseline specimens dispersed throughout each build. A total of 110 specimens were built, 10 were determined to have been impacted by a short-feed event during the build process and their data was discarded. Of the remaining 100 components, one failed an interquartile range outlier test for ultimate flexural strength and was also discarded. The specimens had a nominal size of 65mm x 8.5mm x 3.5mm and were tested in accordance with the ASTM D790 testing standard [13]. They were tested in an orientation such that the crosshead motion was normal to the layers, placing the maximum shear plane parallel with the layers.The next test consisted of two builds of miniature I-beam specimens, one using the ALC technique and the other serving as a baseline for comparison. Each build consisted of 49 identical components whose dimensions are given in Figure 5.

Each build consisted of 49 identical components whose dimensions are given in Figure 5. These components are not overly complex, but they do have step changes in geometry that stress the laser control system more than the simple geometry of the flexure specimens. Since there is not an appropriate mechanical testing standard1677...... ~. •••••• ••••••• ••••• • ••••••• ••••••for these components, they were only evaluated for improvement in post-sintering temperature uniformity.a)b)I"•=t.· l ,Omm. . . . . x • •Figure 5: Dimensions of (a) SLS I-beam test part and (b) their arrangement within the build chamberResultsThe temperature results from the flexure specimens showed a large improvement in uniformity of post-sintering temperatures. To quantify the effects of using the automated laser control system, the average post-sintering temperature of each component was measured on each layer.

To quantify the effects of using the automated laser control system, the average post-sintering temperature of each component was measured on each layer. All components built using the ALC system were compared to determine the standard deviation of post-sintering temperatures on each layer. This was repeated for the fixed power components to provide a baseline measurement. The results are given in Figure 6 and show that the baseline standard deviation is approximately 1.4 °C and constant throughout the build. The ALC components start out around the same level of uniformity, but as the build progresses and the surrogate models collect more data, their prediction capabilities improve and the post-sintering temperature standard deviation is reduced to only 0.5 °C.

Overall, the ALC system improved temperature uniformity by 57%.Post Sintering Temperature STD1.61.4I==Laser ControlledFixed Power1.2)C(n o i t a v e D d r a d n a t Si10.80.60.405101520253035Layer NumberFigure 6: Temperature results from flexure specimens1678The flexure specimens were destructively tested and the ultimate flexural strengths for the two groups were binned and fit with a Gaussian distribution. The two distributions were normalized by their own mean so that a direct comparison of standard deviation could be made. These results are presented in Figure 7(a) and show that the standard deviation of ultimate flexural strengths was reduced by 45% when using the automated laser control system.

These results are presented in Figure 7(a) and show that the standard deviation of ultimate flexural strengths was reduced by 45% when using the automated laser control system. The ultimate flexural strengths are also plotted against their respective post-sintering temperature in Figure 7(b) to highlight how the ALC system drastically reduced the ranges of the strength and post-sintering temperature values, despite having a similar range of pre-sintering temperatures.a)0.050.045Automated Laser Control vs Fixed Power Distribution1-Fixed PowerLaser Controlledb)9080Ultimate Flexural Strength vs Post-Sintering Temp•Laser ControlledFixed Power0.04= 45.28%70y t i l ib a b o r P0.0350.030.0250.020.0150.010.005) a P Mhg n e r tSa r u x e F e t a mli t lU6050403020•.

, •• • ••• • • •• • • • • ••• • •0 -604020020406010191192193194195196UFS DistributionPost-Sintering Temperature (C)Figure 7: Flexural strength results for Automated Laser Control and baseline components shown (a) as normalized probability distribution and (b) against post-sintering temperatureThe I-beam specimens were evaluated for temperature uniformity in the same manner as the flexure specimens. The results, shown in Figure 8, show that the average post-sintering temperature standard deviation was reduced by 45%. The ALC results show an unexplained dip in post-sintering temperature on layer 82 that caused the standard deviation of post-sintering temperatures to be slightly worse than that of the baseline test, though the range was still improved. It is unclear what caused this temperature drop, but the ALC system self-corrected and quickly brought the post-sintering temperature standard deviation back within the desired range.

It is unclear what caused this temperature drop, but the ALC system self-corrected and quickly brought the post-sintering temperature standard deviation back within the desired range. Based on the strength improvements measured by the flexure specimens, it is expected that the strength uniformity of the I-beam specimens was also greatly improved.1679197198Automated Laser Control Improvement80.00%60.00%t n e m e v o r p m40.00%20.00%It n e c r e P0.00%20.00%010203040506070809010040.00%60.00%LayerFigure 8: Improvements in I-beam post-sintering temperature uniformity when using Automated Laser ControlConclusionThis paper covered the implementation of a novel laser control technique for selective laser sintering. The proposed control system was evaluated by comparing the uniformity of components it produced with those produced using a constant laser power.

The proposed control system was evaluated by comparing the uniformity of components it produced with those produced using a constant laser power. The components showed up to 57% improvement in uniformity of post-sintering temperatures and up to 45% improvement in strength uniformity. Adoption of this system on a commercial machine could lead to more uniform components, a reduced scrap rate, and improved confidence.There are a few avenues of future work that may increase the impact of this laser control system. First, the desired temperature value used for testing was found empirically through destructive testing of previously built components. A more sophisticated approach might reveal more optimized values that would further improve the strength uniformity improvements. Another area for future work is to test the system on other materials. So far, all testing has taken place on nylon 12, but the control system should be directly applicable to any SLS material.

Another area for future work is to test the system on other materials. So far, all testing has taken place on nylon 12, but the control system should be directly applicable to any SLS material. Evaluating the performance of the control system on a variety of materials could improve confidence and utility.1680RangeSTDWorks Cited[1] B. Caulfield, P. E. McHugh and S. Lohfeld, "Dependence of mechanical properties of polyamide components on build parameters in the SLS process," J of Mater Process Technol, vol. 182, pp. 477-488, 2007.[2] W. W. Wroe, J. Gladstone, T. Phillips, S. Fish, J. Beaman and A. McElroy, "In-situ thermal image correlation with mechanical properties of nylon-12 in SLS," Rapid Prototyp J, vol. 22, no. 5, pp. 794-800, 2016.[3] S. Negi, S. Dhiman and R. K. Sharma, "Determining the effect of sintering conditions on mechanical properties of laser sintered glass filled polyamide parts using RSM," Mater, vol. 68, pp. 205-218, 2015.[4] I. Gibson and S.

Fulcher, Real-time Pyrometer Feedback and Control in Metal Powder Bed Fusion,Austin: paper presented to Solid Freeform Fabrication Symposium, 2018.[8] S. K. Everton, M. Hirsch, P. Stravroulakis, R. K. Leach and A. T. Clare, "Review of in-situ process monitoring and in-situ metrology for metal additive manufacturing," Mater and Des, vol. 95, no. 2, pp. 431-445, 2016.[9] J.-P. Kruth, P. Mercelis, J. V. Vaerenbergh and T. Craeghs, "Feedback control of Selective Laser Melting," in Proceedings of the 15th International Symposium on Electromachining, 2007.[10] M. Grasso and B. M. Colosimo, "Process defects an in situ monitoring methods in metal powder bed fusion: a review," Measurement Science and Technology, vol. 28, no. 4, 2017.[11] T. Craeghs, C. Stijn, E. Yasa and J.-P. Kruth, "Online Quality Control of Selective LaserMelting," in Solid Freeform Fabr Symp, Austin, 2011.[12] D.

28, no. 4, 2017.[11] T. Craeghs, C. Stijn, E. Yasa and J.-P. Kruth, "Online Quality Control of Selective LaserMelting," in Solid Freeform Fabr Symp, Austin, 2011.[12] D. Rowell, "Introduction to Recursive-Least-Squares (RLS) Adaptive Filters," in SignalProcessing - Continuous and Discrete, Massachusetts, MIT, 2008.1681.

 Evaluation of Cross-Sectional Image Generation Schemes in the Selective Inhibition Sintering (SIS) ProcessMatthew Petros*, Payman Torabi*, Behrokh Khoshnevis*Daniel J. Epstein Department of Industrial and Systems Engineering, University of Southern California, Los Angeles, CA 90089Abstract Selective Inhibition Sintering of metal alloys (SIS-metal) has been proven effective in the additive manufacture (AM) of low resolution bronze parts. The use of high precision inkjet print heads represents a significant advancement in the SIS-metal process. The fabrication of complex three-dimensional metallic parts requires SIS-metal compatible, cross-sectional image processing based on the part boundary profile. Thus, three candidate layer-processing approaches were identified and validated for rudimentary geometries. These approaches were identified from previous research as well as preliminary investigations.

These approaches were identified from previous research as well as preliminary investigations. The validation criteria is based upon maintaining part integrity, the amount of powder waste produced, processing time, the ability to handle various part geometries, and the ease of access to inhibited regions. Results are discussed for deploying the three candidate layer processing approaches for rudimentary shapes, and a preliminary evaluation is presented for their use on more complex geometries.Introduction Additive manufacturing is a term used for the direct manufacture of parts from digital models. AM works by adding material in thin layers from the bottom up, with each layer representing a cross-section of the part [1]. There are various competing technologies in the AM of metal parts such as Direct Metal Laser Sintering (DMLS), Laser Engineered Net Shaping (LENS), Electron Beam Melting (EBM), Selective Laser Melting (SLM) and selective adhesive deposition [2].

There are four common steps to all of these AM technologies:1) A digital model of a part is constructed using one of the various CAD packages(SolidWorks, Alibre, Google SketchUP, and many more).2) The model is imported into a slicing software which generates the profile and/or tool pathinstructions for each layer of the part.3) A file containing the profile and tool path instructions is transferred to the AM machine,which subsequently builds the part one layer at a time.4) The part is removed and post-processed to achieve desired aesthetic and/or mechanicalproperties.Selective Inhibition Sintering Traditional research in powder sintering has mainly focused on enhancing and speeding the sintering process. Similarly, the metal AM processes mentioned above selectively sinter or fuse particles within each layer’s part cross-section. In contrast, the Selective Inhibition Sintering (SIS) process, developed by Dr.

In contrast, the Selective Inhibition Sintering (SIS) process, developed by Dr. Berok Khoshnevis, is based upon the retardation of sintering [2, 3, 4, 5]. The core concept behind the SIS process is the prevention of selected regions of each powder layer from sintering; this is accomplished by treating the regions external to the part in each layer with a sintering inhibitor. In this sense, the SIS process can be considered an inverse to traditional metal AM processes [6].In the current metal process, a commercial piezoelectric print head is utilized to deposit a liquid chemical solution at the periphery of the part for each layer. Once all of the layers have been completed, the entire part is removed from the machine and bulk sintered in a conventional sintering furnace. The inhibitor deposited at the part's boundary decomposes into hard particles that retard the sintering process.

The inhibitor deposited at the part's boundary decomposes into hard particles that retard the sintering process. The particles in this region are prevented from fusing, allowing for removal of inhibited boundary sections and revealing of the completed part. Figure 1 is an illustration of the SIS-metal process.Figure 1: The SIS-metal processIntroduction to Machine Path Planning Each AM machine needs a machine path generated specifically for that machine’s respective process. In all AM processes, with the exception of SIS, the machine path generated traverses the internal cross-section of the part. For example, the standard scanning strategy for the production of parts using SLM first scans the boundary contour and then moves inward to "hatch" the internal cross section for that layer. Various process planning techniques have been developed to accommodate different AM technologies [7].

Various process planning techniques have been developed to accommodate different AM technologies [7]. Factors such as hatch spacing and degree of orientation are important parameters for determining the final mechanical properties of the part [8]. In addition to these parameters, the electron or laser beam speed and layer thickness can have significant effects on part shrinkage and the final size of the part [9].Previous hatching schemes in SIS, however, had a reverse hatching scheme. Asiabanpour et al. utilized a single nozzle solenoid valve to produce plastic SIS parts by depositing inhibitor solution external to the part periphery [10]. This was done in a raster printing scheme. In 2012, Yoozbashizadeh utilized the same path planning for the successful production of SIS-metal parts [11].

This was done in a raster printing scheme. In 2012, Yoozbashizadeh utilized the same path planning for the successful production of SIS-metal parts [11]. In addition to exploring the use of this hatching scheme, two new hatching schemes are considered for deployment in the high-resolution metal-SIS test bed.Experimental Setup and ProceduresHardware The metal SIS beta prototype machine (Figure 2) was developed at the Center for Rapid Automated Fabrication Technologies (CRAFT) and was utilized for the purposes of this research. The system consists of a three-axis motion system which uses a commercial piezoelectric Epson Workforce 30 print head [12, 13].Figure 2: Metal SIS prototype machine.The inhibitor solution is deposited onto the powder bed by activating two color channels of the printhead simultaneously. These two channels collectively account for one hundred and twenty nozzles.

These two channels collectively account for one hundred and twenty nozzles. Two colors were used in order to maximize the amount of fluid deposited per pass of the print head. The inhibitor chosen for this research was common table sugar, or sucrose (C12H22O11). The inhibitor solution consisted of sucrose dissolved in water and an organic surfactant. Building upon the research of Khoshnevis et al. [6], the metal powder used was a fully alloyed bronze with chemical composition shown in Table 1. The powder particle size has a distribution of 98.7% -325mesh, 1.3% -200/+325mesh as reported by the manufacturer.Table 1: Chemical composition of bronze as reported by manufacturerCopper90%Tin10.00%Lead0.025%Zinc0.04%Iron0.058%Phosphorous0.085%Bronze Chemical CompositionUpon completion of inhibitor deposition, green part was sintered in a programmable, hightemperature Neytech Centurion Qex porcelain furnace. Sintering takes place in a vacuum environment.

Sintering takes place in a vacuum environment. The sintering profile can be seen in Figure 3.01002003004005006007008009000100200300400500Temperature [oC]Time [min]Bronze Sintering ProfileFigure 3: Sintering profile used in experimentsAfter sintering, the part is ready for post-processing. Separation or removal of inhibited regions was accomplished using a handheld rotary tool with a wire brush attachment typically used for polishing. A Dremel, Model 3000, was used for the purposes of these experiments.Software A simple Graphical User Interface (GUI) was developed in Visual C# for ease of implementation of various deposition hatching schemes. Parts were first designed and saved into STL format in SolidWorks CAD software. MeshLab was then used to view and manipulate the STL files when necessary (Figure 4.a&b). Based on the aforementioned research of Asiabanpour et al.

MeshLab was then used to view and manipulate the STL files when necessary (Figure 4.a&b). Based on the aforementioned research of Asiabanpour et al. and literature review of other AM slicing approaches, the layered images (Figure 4.c) were constructed using a predetermined uniform layer thickness for a given binary STL file format. The layer thickness was determined by the minimum spreadable layer thickness for bronze powder, 120 µm.(b) Figure 4: (a) SolidWorks model of simple shapes part file (b) MeshLab view of the associated STL file (c) Sliced image of the STL file(a)(c)The cross-sectional profiles were obtained from the intersection between the object and regularly spaced planar faces perpendicular to the z direction. The triangles were first sorted by theirrespective vertex z values in ascending order.

The triangles were first sorted by theirrespective vertex z values in ascending order. For each layer, two intersections were calculated per triangle to define the cross-sectional profile at a given z height (Figure 5).Figure 5: Illustration of two intersection points (red) created by each triangle intersecting a given z planeLoops in a given cross section were differentiated by a simple algorithm. Each triangle produced a pair of intersections, [x1, y1] and [x2, y2], which were stored in an irregular sequence. The pairs were first sorted by checking against other pairs in the layer. The sorting process started from the first pair of intersections in the array and searched for an adjacent pair. Once the adjacent pair was found, the next adjacent pair was searched for. This was continued until the loop closed upon itself.

Once the adjacent pair was found, the next adjacent pair was searched for. This was continued until the loop closed upon itself. If there were points unaccounted for in the same layer, the process would begin again to identify the next loop until all points were accounted for.Each loop was then ranked using the ray casting method to distinguish between hole and part boundaries; this process determines which side of the boundary will be marked for inhibitor deposition. For example, if the loop represents a part boundary, the area external to the loop will be marked for deposition. If the loop represents the boundary of a hole, the area internal to the loop will be marked for deposition (Figure 6).(b) Figure 6: (a) Digital model of a washer (b) Sliced and colored image of a layer cross-section.

Yellow represents part, blue represents the filled (printed) region, and red represents the original part boundary location (c) Final sliced image.(a)(c)An even rank (0, 2, 4…) signifies part boundaries and an odd rank signifies hole boundaries. From these calculations, the images were generated and stored (Figure 6c). For the purposes of this study, images were manually edited depending on the hatching scheme being used. Finally, a print command was sent to the Epson Workforce 30 to begin inhibitor deposition.Experimental Objectives Based on the results of preliminary experiments, three hatching schemes were assessed for compatibility with the SIS process. The objectives of these hatching schemes were to:A. Minimize powder waste B. Maintain part integrity C. Minimize process time D. Evaluate the ability to handle complex part geometriesHatching Scheme A (HSA) Previous research suggested a hatching scheme in which an inhibited base is printed in the first layer.

Evaluate the ability to handle complex part geometriesHatching Scheme A (HSA) Previous research suggested a hatching scheme in which an inhibited base is printed in the first layer. The entire part negative is filled with inhibitor solution for each successive layer afterward (Figure 3c). This hatching scheme is simple to implement and allows for ease of post-processing in that there is direct access to inhibited regions for removal after sintering. However, the large amount of waste is a significant drawback to this hatching scheme. There is currently no cost- effective method for recycling metal powder that has been contaminated by inhibitor solution. In the case of HSA, the entire negative region is contaminated.The experiment was designed with rudimentary shapes as a proof of concept. The experiment consisted of extruded square, circle, pentagon, and triangle shapes. The same image was repeated over ten layers to create rudimentary shape extrusions.

The experiment consisted of extruded square, circle, pentagon, and triangle shapes. The same image was repeated over ten layers to create rudimentary shape extrusions. Figure 7a illustrates the results of HSA after sintering.(a)(b)(c)(d)Figure 7: (a) Rudimentary shapes after sintering. The darker colored regions represent the inhibited region (b) Close-up view of flooding in the square shape (c) Front side of final parts after separation from the inhibited region (d) Back side of final partsThroughout the printing of HSA, it was difficult to control excess fluid deposition. Each layer was visually monitored for signs of flooding. The amount of deposition was adjusted for each layer in an attempt to mitigate part flooding; nonetheless, some layers were flooded in the process. This is evident in Figure 7b, a close-up view of the square shape. Figure 7cd represent the result of the final part after removal of the inhibited regions.

This is evident in Figure 7b, a close-up view of the square shape. Figure 7cd represent the result of the final part after removal of the inhibited regions. The final parts were deformed and exhibited evidence of cracking. This is believed to be a result of excess fluid deposition that penetrated into the part periphery.As a result of the excess fluid deposited, another drawback was discovered. The excess fluid from layers beneath the surface visibly erupted during the sintering cycle. This is evidenced by the powder ejected in the upper regions of the part. It is believed that the excess water wasn't thoroughly removed during the furnace pre-heat. Consequently, the fluid evaporated rapidly at higher temperatures and was trapped by the liquid and powder above, which eventually led to the eruption.Hatching Scheme B (HSB) Based on the results from HSA and differential shrinkage observed by Yoozbashizadeh [11], HSB called for a reduction in the area of inhibited region.

This serves two purposes: reducing the powder waste and hampering the potential for cracking. Thus, instead of printing the entire part negative region as in Figure 3c (blue), a thinner boundary profile was designed (Figure 8a). In HSB, an inhibited base was again printed in the first layer. The experiment consisted of an extruded square, circle, pentagon, and triangle with 500 µm boundaries surrounding each shape. A 2 mm thick square perimeter was printed to provide a protective bounding box for ease of transport of the green part. The same image was repeated over ten layers to obtain the desired shape extrusions.

The same image was repeated over ten layers to obtain the desired shape extrusions. Figure 8b illustrates the results immediately after sintering, and Figure 8cd illustrate the results after removal of the inhibited regions.(a)(b)(c)(d)Figure 8: (a) Rudimentary shape design with 480 µm borders and a 2 mm thick printed perimeter (b) Results immediately after sintering (c) Front side of final parts after separation from the inhibited region (d) Back side of final partsUpon removal from the furnace, the results showed significant visible improvement. The triangle, circle, and pentagon shapes were successfully extracted. However, the square shape exhibited internal cracking. Cracks were observed at the corners of the printed bounding box, as well as within the bounding box (Figure 8b).

However, the square shape exhibited internal cracking. Cracks were observed at the corners of the printed bounding box, as well as within the bounding box (Figure 8b). Slight differential shrinkage was observed at the corners of the bounding box and the triangular part.Similar to HSA, the extracted rudimentary shapes exhibited signs of excess fluid deposition. This was evidenced during the inhibitor removal. While removing inhibited regions from the back side of the parts, material from within the part itself was easily removed. This signifies that the liquid inhibitor once again penetrated into part.Hatching Scheme C (HSC) The results from HSB showed visual improvement for reducing the amount of inhibitor solution deposited. Three of the four original shapes were extracted successfully. However, the integrity of the parts was not maintained. Thus, the HSC design was based on the minimum amount of inhibitor solution allowable for the designed shapes. Torabi, et al.

However, the integrity of the parts was not maintained. Thus, the HSC design was based on the minimum amount of inhibitor solution allowable for the designed shapes. Torabi, et al. experimentally determined the minimum line thickness for boundary inhibition to be approximately 200 microns. For the SIS process, this is a minimum gap test because the thinnest separable printed line becomes a gap in the finished part after abrasive blasting. Figure 9a represents the sliced image design for the thinnest separable part boundary. The printed perimeter surrounding the shapes in HSB was removed to alleviate excess deformation and stresses induced by water evaporation and shrinkage. In addition, the inhibited base layer was removed in HSC. Instead, the design image (Figure 9a) was printed as the first layer.

In addition, the inhibited base layer was removed in HSC. Instead, the design image (Figure 9a) was printed as the first layer. Figure 9b illustrates the results of the HSC experiment immediately after sintering, and Figure 9cd represent the final part results.(a)(b)(c)(d)Figure 9: (a) Rudimentary shape design with 200 µm shape borders (b) Results of the HSC experiment immediately after sintering (c) Final part resultsThe results of HSC again showed significant improvement. There was no visible evidence of cracking within the rudimentary shapes nor external to the parts. Differential shrinkage was not as obvious with the thin border design. In addition, there was no evidence of over penetration of inhibitor within the part periphery. In other words, all deposited inhibitor remained outside of the part as intended.Discussion The parameters introduced as experimental objectives were evaluated in Table 2.

In other words, all deposited inhibitor remained outside of the part as intended.Discussion The parameters introduced as experimental objectives were evaluated in Table 2. The ability of a printing scheme to address complex geometries is dependent upon physical access to the inhibited regions after sintering. In this regard, while HSA can theoretically handle the most complex geometries, the amount of powder waste was high and the integrity of the part was very poor. The size of the shapes as well as the verticality of the walls of shapes was a good indicator of shape integrity. HSB had significantly lower powder waste when compared with HSA, but the part integrity was still poor. The integrity was, however, an improvement over HSA. The size of the shapes was closer to the designed size, but the walls were not vertical. In addition, the square shape was cracked internally. HSC produced the lowest amount of powder waste of all, and theparts maintained a high integrity.

In addition, the square shape was cracked internally. HSC produced the lowest amount of powder waste of all, and theparts maintained a high integrity. The experiment resulted in shapes that were close to their nominal design size and the walls were vertical.Table 2: Evaluation Criteria for the three hatching schemes.Printed Area [mm2]Percent WasteGeometric ComplexityPrinting Time [sec/layer]Access to Inhibited RegionsPart IntegrityHSA146780%HIGH19.7HIGHLOWHSB45925%LOW19.6LOWLOWHSC633%LOW16.3LOWHIGHConclusion and Future Work Hatching Scheme C presented the most promise for the fabrication of high-quality metallic parts using the new Selective Inhibition Sintering metal prototype machine. However, using HSC only allows for the fabrication of limited part geometries. Complex three-dimensional shapes could not be physically extracted after sintering due to inaccessible inhibited regions.

(1289076858).[12] Ink Jet Printers Technical Brief. (2007, May 1). Epson[13] Epson Micro Piezo Print Head Technology. (2010, January 1). Epson Exceed Your Vision..

aluation of Direct Diode Laser Deposited Stainless Steel 316L on 4340 Steel Substrate for Aircraft Landing Gear Application11121Tian Fu, Todd E. Sparks, Frank Liou, Joseph Newkirk 1, Zhiqiang Fan3, Syamala R. PulugurthaJianzhong Ruan, Hsin-Nan Chou1Department of Mechanical and Aerospace Engineering, Missouri University of Science & Technology2Department of Metallurgical Engineering, Missouri University of Science & Technology3Boeing Phantom WorksCorresponding Address: 292B Toomey Hall, 400 W. 13th St., Rolla, MO 65409-0050300M steel is used extensively for aircraft landing gears because of its high strength, ductility and toughness. However, like other high-strength steels, 300M steel is vulnerable to corrosion fatigue and stress corrosion cracking, which can lead to catastrophic consequences in the landing gear. Stainless steels offer a combination of corrosion, wear, and fatigue properties.

Stainless steels offer a combination of corrosion, wear, and fatigue properties. But for an aircraft landing gear application a higher surface hardness is required. A laser cladding process with fast heating and cooling rates can improve the surface hardness. AISI 4340 steel is used as a lower cost alternative to 300M due to its similar composition. In this study, the influence of laser cladding process parameters, shield gas, and composition of the deposition and dilution zone has been investigated. The microstructures and composition analysis were evaluated by Scanning Electron Microscopy (SEM) and Optical Microscopy. The deposition hardness varies from 330HV to 600HV.1. IntroductionHard chrome plating is a technique that has been the “the gold standard” in commercialproduction for more than 50 years. It is also a critical process that is used for applying hard coatings to a variety of aircraft components in manufacturing operations.

It is also a critical process that is used for applying hard coatings to a variety of aircraft components in manufacturing operations. However, hard chromium plating baths contain chromic acid in which the chromium is in the hexavalent state. Hexavalent chromium (hex-Cr) is a known carcinogen having a level of toxicity greater than arsenic or cadmium. So the health risks associated with hard chromium have driven the engineering community in both the military and civilian aerospace sectors, as well as many industrial sectors, to seek alternative coating materials, as well as more cost effective methods to apply them [1].Among the most promising alternatives are laser cladding coatings. Laser claddingcoatings are becoming more industrially viable as lasers have improved dramatically in the past few years, with the production of high power Nd-YAG, CO2, and diode lasers that are robust and designed for industrial use.

The method can be used to weld onto the surface any compatible alloy that can be delivered in wire or powder form. In general the method requires reasonable4702,access for the laser beam and the feed material, with the laser striking the surface at roughly normal incidence [2].The cladding process is faster than conventional chrome plating. Typical claddingcoatings are applied in one hour, versus 24 hours or more required for hard chrome plating. Another advantage of this process is coating adhesion is excellent due to the metallurgical adhesion, rather than the mechanical bond which High Velocity Oxygen Fuel(HVOF) process produces [3].Laser cladding tends to be used for high temperature materials. For example, GE Aircraft Engines uses this method for coating the tips of some hot section turbine blades. It is not easy to set up the correct processing parameters, but once they are determined the method can be very robust.

It is not easy to set up the correct processing parameters, but once they are determined the method can be very robust. This process can produce porosity free deposit more readily than HVOF, which usually produces 2%-5% porosity [3].The Laser Aided Manufacturing Processes (LAMP) system which was developed at Missouri S&T mainly consists of a 5-axis CNC machining center, a 1.0 KW Diode laser, a powder feeder, and a real–time control system from National Instruments. During deposition, the substrate is fixtured on a 5-axis CNC. The nozzle through which the laser and metal powder is transmitted is fixed to the Z-axis of the CNC. The laser is focused on a small area of the substrate and creates a molten pool, and the metal powder is delivered by the powder feeder system into the molten pool to create the deposition. The X, Y and Z table positions and velocities are regulated via the CNC machining center controller according to the program generated from the CAD model [4].

The X, Y and Z table positions and velocities are regulated via the CNC machining center controller according to the program generated from the CAD model [4]. This hybrid repair system employs a 5-axis positioning system which includes 3 linear axes and 2 rotating axes. The advantage of using this system instead of using the conventional 3-axis positioning system is that it does not require support material to build overhang features for 3D parts. This capability allows both the deposition and machining in a single set-up for a part even with intricate or hidden features. Figure 1 shows the Direct Laser Deposition process.Fig. 1: Laser Aided Manufacturing Process deposition system4712. Experimental procedureIn the present study, gas atomized 316L stainless steel (purchased from Carpenter Technology) powders of mesh sizes -70+325 was deposited on AISI 4340 steel (Purchased from Rose Metal Products, Inc). Table 1 shows the composition of 316L stainless steel and AISI 4340 steel.

%) 316L steel Composition (wt. %)C0.15- 0.350.2- 0.3 Bal.1.65- 2.000.7- 0.90.38- 0.430.0350.0350.6-0.816-10- 14%<2%<1%<0.03%Bal.<0.03%18.5% 2-3%<0.045%A cross-section of the deposit and the substrate was prepared to observe themicrostructure under the optical microscope (Nikon) and scanning electron microscope (Hitachi, model S570). The samples were etched by a molybdic acid reagent, composed of 100 ml each of HNO3, HCL, and distilled water with 3 grams molybdic acid. This is a swab etch: A cotton ball was placed in a beaker of etchant and tongs were used to swab the sample surface for approximately 10 seconds. The secondary electron images and EDX data were captured at 10kV accelerating voltage and 15mm working distance [5]. Microhardness of the fabricated layer (from the top surface to the bottom of substrate) was measured using a Vicker’s microhardness tester (Struers, Duramin 5) with a 100g applied load.a. Argon shield gas sample b.

Argon shield gas sample b. Nitrogen shield gas sampleFig. 2 AISI 316L stainless steel powder deposited on AISI 4340 steel substrate with argon and nitrogen4723. Results and discussionsHomogeneous and defect free 316L stainless steel coating microstructure has been abstained. Fig. 3 shows the cross-sectional microstructure of laser assisted fabricated 316L stainless steel processed with a laser power of 850W, scan speed of 13.7 ipm and a powder feed rate of 4.5 g/min in both deposition and dilution zone. The coating is dense, adherent and crack- free. The microstructure is mostly cellular, with an average grain size of 10µm. An optical microscopy study shows that the coating microstructure is highly dependent on laser cladding process parameters and under a narrow range of processing parameters a defect free and homogeneous microstructure can be obtained [6].Fig.

3 Homogeneous microstructure of laser fabricated 316L stainless steel on 4340 steel substrateThe microhardness analysis of the fabricated product showed that the average microhardness varied from 330 to 600HV across all the samples in this study. The hardness variation was observed in both the deposit and the dilution area. The hardness measured is significantly higher than the microhardness values reported in previous work with similar materials and process conditions where the microhardness values varied from 150 to 280HV [6]. Fig. 4 illustrates the microhardness trend from the top down (opposite growth direction) for laser assisted fabricated 316L stainless steel on AISI 4340 steel processed with a laser power of 760 W, scan speed of 10 inch/min, and a power flow rate of 3.6 g/min. The behavior shown in Fig. 4 is typical for all samples generated in this study.473Fig.

The behavior shown in Fig. 4 is typical for all samples generated in this study.473Fig. 4 The trend of microhardness of laser deposited 316L stainless steel on AISI 4340 steel along the direction from the top down (Process parameters: laser power of 760 W, scan speed of 10 inch/min; power flow rate of 3.6 g/min)Fig. 5 is a quantitative analysis as a function of depth. From Fig. 4 and Fig. 5 we can conclude that the elemental composition of the deposition and the dilution zone are almost the same. Also, the hardness of the deposition and dilution zone is similar. At the bottom of the dilution zone, we can see the sudden composition change as well as the hardness change. The hardness change is because the composition of the dilution zone and heat-affected zone differs. The hardness of the heat-affected zone can reach as high as 600HV, comparing the annealed AISI 4340 steel substrate hardness from 250 HV to 300 HV.

The hardness of the heat-affected zone can reach as high as 600HV, comparing the annealed AISI 4340 steel substrate hardness from 250 HV to 300 HV. In the present study, no significant microhardness as well as microstructural difference between shield gas nitrogen and argon has been found, thus it should be cost effective to use nitrogen as shield gas.474Fig. 5 Quantitative analysis and as a function of depth for laser deposited 316L stainless steel on AISI 4340 steel from the top surface to the bottom of the substrate (Process parameters: laser power of 760W, scan speed of 10 inch/min, power flow rate of 3.6 g/min)4. Summary and conclusionsIn this paper, the microhardness and EDX analysis of laser assisted fabricated AISI 316L stainless steel on the 4340 steel substrate has been studied with the following process condition: laser powder of 700-1000W, scan speed of 10-20 ipm and powder flow rate of 3- 6 g/min.From the results, the following conclusions could be drawn:1.

The microhardness of the fabricated layer was significantly improved to as high as 330–600 HV as compared to 150 VHN when conventionally processed. The improved microhardness is attributed to grain refinement achieved during laser processing as well as the fast cooling rate during the laser cladding process. 2. No significant microhardness or microstructural difference between shield gas nitrogen and argon has been found, thus it will be cost effective to use nitrogen as shield gas. 3. The elemental composition of the deposition and the dilution zone are almost the same. Also, the hardness of the deposition and dilution zone is similar.4755. AcknowledgementsThis research was supported by the National Science Foundation Grant Number DMI-9871185, the grant from the U.S. Air Force Research Laboratory contract # FA8650-04-C-5704, and Missouri S&T Intelligent Systems Center. Their support is greatly appreciated.6. References[1] B. D. Sartwell, P. M. Natishan, I. L.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEvaluation of Early Fatigue Damage Detection in Additively Manufactured AlSi10MgSusheel Dharmadhikari1 and Amrita Basak1, *1Department of Mechanical Engineering, The Pennsylvania State University, University Park, PA 16802Corresponding Author: aub1526@psu.eduAbstractThe article presents two distinct measures for fatigue damage detection in additively manufactured AlSi10Mg specimens with a one-sided V-notch. The specimens are fabricated on a ProX-320 laser powder bed fusion equipment with recycled AlSi10Mg powders using vetted process parameters as suggested by 3D Systems. The process of fatigue damage evolution is monitored using two heterogeneous sensing techniques, namely, the force-displacement sensors and a confocal microscope.

The process of fatigue damage evolution is monitored using two heterogeneous sensing techniques, namely, the force-displacement sensors and a confocal microscope. The force and displacement sensors are embedded in the fatigue testing apparatus to capture the global effects of the stress-strain behavior of the specimens; however, it provides no information about the local damage near the notch. The force-displacement time-series data, which shows a hysteresis-like behavior, is calibrated using a confocal microscope focused inside the notch of the specimen so that the onset of fatigue crack initiation can be detected at a crack opening displacement (COD) of ~10 microns. Using the force-displacement data, the energy dissipation rate and the material stiffness per cycle are computed. The results show a detection accuracy of 96.25% and 90.84% for the energy dissipation rate and material stiffness per cycle, respectively.

The results show a detection accuracy of 96.25% and 90.84% for the energy dissipation rate and material stiffness per cycle, respectively. In conclusion, the paper establishes two successful predictors for fatigue damage detection in additively manufactured AlSi10Mg specimens.Keywords Fatigue Failure Detection; Laser Powder Bed Fusion; Force-Displacement Sensors; Confocal Microscope1. IntroductionWhile over the past few decades, additive manufacturing (AM) has generated a tremendous amount of excitement within the materials and manufacturing community, sub-par mechanical properties (e.g., fatigue) continue to challenge the researchers across industry and academia [1]. A considerable volume of research exists with regards to online real-time fatigue damage detection for traditionally fabricated metallic materials via different sensing technologies such as ultrasonics [2], acoustic emissions [3], eddy currents [4], etc.

The fundamental working principle of such sensing techniques are based on three cardinal pillars: (i) acquisition of the sensor data in the form of a time-series; (ii) classification (e.g., healthy and cracked) of the time-series data using a secondary sensor (frequently an imaging one such as a digital or an optical microscope); (iii) application of time-series pattern recognition techniques to pinpoint the damage solely based on time-series without any requirement of the secondary sensor as it would be unrealistic to have an imaging sensor when a component is operating in the field. The efficacy of the method sensors is acid-tested by their capability in detecting the length-scale of fatigue cracks [4] which consist of a805range of length scales starting from microns to millimeters [5]. The state-of-the-art on traditionally manufactured components shows successful damage detections when the crack opening displacement (COD) is in the millimeter scale length [6].

The state-of-the-art on traditionally manufactured components shows successful damage detections when the crack opening displacement (COD) is in the millimeter scale length [6]. COD, as the name suggests, corresponds to the distance between the ends at the base of a crack. A representative notched region of the specimen with a schematic for COD (and crack length) is shown in Fig. 1(a) and (b). Traditionally, the research has been focused towards studying crack length but with recent studies [5], it has been shown that COD is an earlier indicator of crack emergence.

1(a) and (b). Traditionally, the research has been focused towards studying crack length but with recent studies [5], it has been shown that COD is an earlier indicator of crack emergence. Irrespective of whether the component is traditionally or additively fabricated, an important research direction in this area is, therefore, to enable the detection of fatigue damages in the micron-scale COD [5].Crack LengthCrack opening displacement(COD)(a)(b)Figure 1: (a) An overview of the notched region of a specimen with a crack, (b) Schematic highlighting the crack opening displacement (COD)To achieve this progression and that too for AM components, the research presented in this article develops a method for fatigue damage detection at the micron-scale COD using an energy dissipation approach [7, 9] where a high-precision confocal microscope is used for data calibration.

Energy dissipation due to material hysteresis has been carried out with strain-controlled experiments for traditionally fabricated metallic specimens to date [7]. This article presents a complementary energy dissipation interpretation for stress-controlled experiments for AM specimens. The study characterizes the force and displacement time-series data during fatigue tests collected using an embedded load cell and a linear variable differential transformer (LVDT), respectively. Using the time-series data, an energy-dissipation-based metric and a stiffness metric are derived and their efficiencies in detecting fatigue damages in AM AlSi10Mg specimens are quantified.2. Experimental2.1. Materials and SpecimenRecycled AlSi10Mg powder obtained from 3D Systems is used in the present study. The composition of the powder as obtained via energy dispersive x-ray spectroscopy (EDS) is: 86.83 wt. % Al, 11.83 wt. % Si, 0.34 wt. % Mg, 0.96 wt. % O, and 0.04 wt. % Fe.

A striped hatch design pattern with hatch spacing 82 μm is used. The specimens used in the tests are in as-built condition without any post-processing.(a)(b)Build direction160.0080.004.00R1220.00Figure 2: (a) Geometry of the one-sided V-notched specimen (dimensions in mm). (b) An image of the actual specimen with an American quarter for scale.2.2. Fatigue Testing ApparatusThe fatigue experiments are performed on an MTS Elastomer 831.10 servo hydraulic setup (Fig. 3(a)) which is equipped with a high-precision confocal microscope. The fatigue test is configured to apply a sinusoidal uniaxial tensile loading with a maximum force of 3 kN and a stress ratio (min. force / max. force) of 0.33 at a frequency of 20 Hz resulting in a maximum nominal stress of 93.75 MPa and a stress amplitude of 31.2 MPa in the specimen. Based on the fatigue strength of AlSi10Mg, these loading conditions ensure a stress-dominated low-cycle fatigue failure in the specimens.

Based on the fatigue strength of AlSi10Mg, these loading conditions ensure a stress-dominated low-cycle fatigue failure in the specimens. The fatigue test is controlled through an automated routine from the Multi-Purpose TestWareTM (MPT) software suite available for the MTS FlexTest® controller. The MTS system can measure the force (F) and displacement (D) responses in the form of a time- series during a test using an embedded load cell and LVDT, respectively. The F-D data, which is acquired through the MPT routine, is captured at 6144 Hz. To calibrate the F-D time-series data for pinpointing the onset of fatigue damage initiation, a high-precision confocal microscope (Bruker Alicona, Graz, Austria) is used. The microscope can detect fatigue damages with a COD of 3 m. A recent study [5] shows that the microscope orientation shown in Fig. 3(b) can detect a fatigue crack earlier. Fig. 3(c) shows the image of the notch base before the testing.

A recent study [5] shows that the microscope orientation shown in Fig. 3(b) can detect a fatigue crack earlier. Fig. 3(c) shows the image of the notch base before the testing. During the test, the MPT routine is programmed such that the periodical loads are stalled every 500 cycles.807This stall is maintained at the maximum load and the period during the stall is used to monitor the base of the notch using the confocal microscope. If a crack is observed during the monitoring, the imaging is performed.Confocal MicroscopeConfocal microscope focus area –notch root(b)(a)ConfocalHigher stress concentrations at the interface between the notch root and side surfaceMoving Stage200 m(c)Figure 3. (a) Customized fatigue testing apparatus showing the specimen, the MTS testing system, the confocal microscope, and the moving stage. (b) A schematic illustrating the focus of the confocal microscope at the base of the notch.

Fig. 5(a) and 5(b) show the F-D time-series data for a representative (here, four) cycles measured through the load cell and LVDT, respectively. The F data shows a comparatively smoother operation due to the controlled application, however, the D data shows some noisy behavior around the peaks which can be attributed to measurement artifacts.dCracked050001000015000No. of load cyclese(a)aHealthy00.10.20.30.40.50.6Displacement (mm)(b)Figure 6: (a) D data during the entire fatigue test for a representative specimen with three indicators correlated with the microscope images in Fig. 4(a), (d), and (e). (b) Hysteresis behavior shown by the F-D time-series data.The sinusoidal nature of the D data, however, is a local behavior as due to the damage accumulation, the D data of the specimen shows a gradual growth in its average magnitude as shown in Fig. 6(a) which is plotted against the number of time steps.

6(a) which is plotted against the number of time steps. Captured at 6144 Hz, the F- D time-series data is rich in the number of data points (i.e., timesteps) with roughly 300 of them per cycle. For example, if a specimen is having a fatigue life of 13000 cycles, the F vs. D curve will have 13000 × 300 = 3.9 million data points in it. The D data (Fig. 6(a)) is classified into healthy and cracked regions based on the synchronous information obtained from the confocal microscope. The gradual increase in the displacement is attributed to the increase in strain and subsequent crack growth in the specimen. The transition from the healthy region to the cracked region corresponds to the timestep where the confocal microscope measures a crack with COD 10 m which is roughly in the short crack regime [5]. Fig. 6(b) shows the hysteresis-like nature by plotting the F-D time series data.

Fig. 6(b) shows the hysteresis-like nature by plotting the F-D time series data. The gradual increase in the area under the curve is attributed to the energy dissipated during the fatigue test. The curve is again classified into the healthy and cracked regions. The disparity between the energy dissipation before and after a crack develops is evident from Fig. 6(b).3.3. Energy Dissipation Rate (EDR)Motivated by the drastic difference in the distribution of the energy dissipation of the healthy and cracked regions as shown in Fig. 6(b), the rate of energy dissipation is hypothesized to contain810the information regarding the fatigue damage. Conventionally, the energy dissipated due to hysteresis is computed with the area under the curve during one cycle [7]. Since this approach is suitable for low cycle fatigue phenomena where the area under one cycle is significant, the approach is deemed to work for the present case as well. Fig.

Since this approach is suitable for low cycle fatigue phenomena where the area under one cycle is significant, the approach is deemed to work for the present case as well. Fig. 7 illustrates the total change in displacement (D, mm) and force (F, N) during N cycles. Using these values, the energy dissipation rate (EDR) is defined as:𝐸𝐷𝑅 =𝛥𝐹 × 𝛥𝐷 𝑁DFFigure 7: Hysteresis behavior for a representative N fatigue cycles showing the total change in force (F) and displacement (D).Healthy2.833.23.43.63.8EDR (N.mm/cycle)2.833.23.43.63.8EDR (N.mm/cycle)Optimal Threshold020406080100% Fatigue LifeCracked020406080100% Fatigue Life2.833.23.43.63.8EDR (N.mm/cycle)HealthyCracked(a)(b)Healthy02040% Fatigue LifeCrackedFigure 8: (a) EDR behavior with respect to the normalized fatigue life.

(b) EDR curve with the optimal threshold computed using receiver operating characteristics (ROC).811Therefore, with this definition, EDR denotes the average loss of energy during N cycles. In this study, N is maintained at 100 indicating that EDR is computed at every 100 cycles. The behavior of EDR during the entire fatigue test is shown in Fig. 8(a). For example, if a specimen has a fatigue life of 13000 cycles, the EDR curve will have 13000/100 = 130 data points in it. Such a volume of data is sufficient for any pattern recognition techniques as reported in the open literature. As shown in Fig. 8(a), the specimen exhibits an EDR of around 2.82 N.mm/cycle at the beginning of the test. Each data point on the curve corresponds to the average energy lost during the 100 preceding cycles. The fatigue crack (with COD of 10 m) detection through the confocal microscope is achieved at 0.7 of the normalized fatigue life.

The fatigue crack (with COD of 10 m) detection through the confocal microscope is achieved at 0.7 of the normalized fatigue life. Shortly after the detection, the EDR tends to show a steady growth which is exponential near the end of the life where a fracture is imminent.Table 1: Confusion matrix for EDR.EDRPredicted HealthyPredicted CrackedTruly Healthy100%0%Truly Cracked7.5%92.5%Such a sensitivity of EDR to the damage detection presents an excellent tool towards fatigue damage detection. Combining this information from EDR curve and machine learning, a classifier is built using an optimal thresholding-based technique using the receiver operating characteristics (ROC) to differentiate between a healthy and a cracked specimen [8]. The ROC computation provides an optimal threshold which acts as the boundary between a healthy state and a cracked state.

The ROC computation provides an optimal threshold which acts as the boundary between a healthy state and a cracked state. During the fatigue test, if the EDR values are below this optimal threshold, then it can be inferred that the specimen is healthy. If the test indicates that the EDR is higher than the threshold, then the specimen can be assumed to be cracked. Fig. 8(b) plots the optimal threshold along with a representative EDR curve. The location where the EDR curve crosses the threshold corresponds to the damage detection. A mean threshold of 2.831 is observed for the EDR curve indicating the robustness of the metric. The accuracy of the classification from the workflow is tabulated through a confusion matrix [8] shown in Table 1.

The accuracy of the classification from the workflow is tabulated through a confusion matrix [8] shown in Table 1. This confusion matrix shows the classifier can identify the healthy and cracked specimens with an accuracy of 100% and 92.5%, respectively.3.4.Material StiffnessIn addition to that, the slope of the F-D time-series data is an important feature as illustrated from Fig. 6(b). The slope at the end of the red region is smaller than the slope at the beginning of the test. This slope corresponds to the material stiffness and the drop in it is attributed to material softening with the increase in fatigue damage [6]. From an energy perspective, the stiffness is an indicator of energy storage in the material.

From an energy perspective, the stiffness is an indicator of energy storage in the material. If F and D denote the total force and total displacement during one cycle, respectively, the stiffness during that cycle is represented as:812𝑆𝑡𝑖𝑓𝑓𝑛𝑒𝑠𝑠 =𝐹 𝐷×160 32×1 1000The numerical multipliers 160 and 32 represent the length and area of the cross-section and are used to convert the stiffness into GPa. The above computation is performed for every cycle and the result is shown in Fig. 9(a).The drastic drop in the material stiffness after the damage detection via the confocal microscope can be clearly observed. Like EDR, the damage detection rules are established for the stiffness curve. Fig. 9(b) shows the optimal threshold for a representative stiffness curve. The threshold shows a value of 71.57 GPa.

Like EDR, the damage detection rules are established for the stiffness curve. Fig. 9(b) shows the optimal threshold for a representative stiffness curve. The threshold shows a value of 71.57 GPa. Using the stiffness curves, an accuracy of 82.54% is observed in detecting the cracked specimens as opposed to 99.14% in detecting healthy specimens as shown in Table 2.(a)(b)020406080100% Fatigue Life2.833.23.43.63.8EDR (N.mm/cycle)HealthyCracked020406080100% Fatigue Life2.833.23.43.63.8EDR (N.mm/cycle)HealthyCrackedFigure 9: (a) Stiffness behavior with respect to the normalized fatigue life. (b) Stiffness curve with the optimal threshold computed using ROC.Table 2: Confusion matrix for stiffness.StiffnessPredicted HealthyPredicted CrackedTruly Healthy99.14%0.86%Truly Cracked17.46%82.54%4. Conclusion and Future WorkThe article presents two distinct metrics to assess fatigue damage detection in AM AlSi10Mg alloys.

Conclusion and Future WorkThe article presents two distinct metrics to assess fatigue damage detection in AM AlSi10Mg alloys. EDR and material stiffness show an accuracy of 96.25% and 90.84%, respectively for813detecting a crack with a COD of 10 m. It is to be noted that the metrics presented in this study are specific to the loading conditions used in this study. A more detailed parametric analysis exploring the dependence of these metrics on loading conditions can expand the application of the proposed analysis to more general scenarios. Moreover, the distribution of the metrics for large number of specimens can also shed light on the consistency of the manufacturing process. In the future, the current work will be extended to study the behaviors over larger number of specimens with the effects of various build directions, specimen geometries, and AM process parameters [9].5.

AcknowledgmentsThe research reported in this paper is supported in part by the Department of Mechanical Engineering at the Pennsylvania State University, University Park, PA 16802. Any opinions, findings, and conclusions in this paper are those of the authors and do not necessarily reflect the views of the supporting institution.6. References[1][2][3][4][5][6][7][8] [9]A. Yadollahi and N. Shamsaei, "Additive manufacturing of fatigue resistant materials: Challenges and opportunities," International Journal of Fatigue, vol. 98, pp. 14-31, 2017. S. Gupta, A. Ray, and E. Keller, "Online fatigue damage monitoring by ultrasonic measurements: A symbolic dynamics approach," International Journal of Fatigue, vol. 29, no. 6, pp. 1100-1114, 2007. N. Eleftheroglou and T. Loutas, "Fatigue damage diagnostics and prognostics of composites utilizing structural health monitoring data and stochastic processes," Structural Health Monitoring, vol. 15, no. 4, pp. 473-488, 2016. J. M.

viewed PaperEVALUATION OF FUNCTIONALLY GRADED LATTICE PROPERTIES OF LASER POWDER BED FUSED STAINLESS STEEL 316LB. B. Ravichander, S.H. Jagdale, S. Theeda, G. KumarDepartment of Mechanical Engineering, The University of Texas at Dallas, Richardson, TX 75080AbstractThe development of metal Additive Manufacturing (AM) techniques, in particular the laser powder bed fusion (LPBF) process, has led to an increase in the innovative design and fabrication of lightweight and complex porous metal structures. Despite the limitations of the LPBF process which limits the geometric accuracy of the porous structures, it eliminates the difficulties presented by conventional manufacturing techniques in the fabrication of highly complex structures. The properties of as-built porous structures depend on the unit cell design and porosity level. These lightweight metal structures have applications in medical and aerospace fields.

The properties of as-built porous structures depend on the unit cell design and porosity level. These lightweight metal structures have applications in medical and aerospace fields. The relationships between the lattice geometry and performance must be determined to successfully implement the functional lattice designs. In this study, functionally graded lattice structures are fabricated from steel using SLM technique and the effect of different lattice types on the manufacturability, density and mechanical properties are investigated.Keywords: Laser Powder Bed Fusion, Functionally Graded Lattice, Porosity, SS316L, Part Density1. INTRODUCTIONStainless steel 316L (316L), is an austenitic steel composed of chromium-nickel- molybdenum alloy [1, 2]. 316L is known for its excellent corrosion resistance, oxidation resistance and low neutron absorption rate [3].

316L is known for its excellent corrosion resistance, oxidation resistance and low neutron absorption rate [3]. 316L is one of the most widely used structural materials because of good formability, mechanical strength, and fracture toughness [4, 5]. In addition to structural applications, 316L is also used in orthopedic implants, petrochemical, and nuclear power plants [3]. The mechanical properties and the corrosion resistance of 316L depend on the processing technique [6]. The geometrical complexity of metal parts in new applications is constantly increasing and the conventional manufacturing techniques like machining, casting, and forging are not suitable to meet the demand [7]. The advances in computer aided design (CAD) have led to new design opportunities like topology optimization that enables reduction in part weight without compromising the performance.

The advances in computer aided design (CAD) have led to new design opportunities like topology optimization that enables reduction in part weight without compromising the performance. Due to the changes in the design and the desired properties, there is a need for the microstructure and mechanical properties to be controlled at various levels. Therefore, AM offers a viable alternative to conventional manufacturing techniques [8, 9].LPBF is a layer-by-layer manufacturing process capable of fabricating highly dense parts with complex geometries [10-14]. The LPBF process uses a recoater to deposit metal powder on the substrate. The powder is then melted and fused with the help of a laser heat source. The process is continued till the entire geometry of the part is fabricated [14-16] . LPBF process has the ability2139to fabricate specimens with varying porosities and complex shapes.

The process is continued till the entire geometry of the part is fabricated [14-16] . LPBF process has the ability2139to fabricate specimens with varying porosities and complex shapes. The LPBF fabricated parts can achieve microstructures and properties comparable to the cast counterparts under optimized processing conditions. The customized microstructures and properties can be achieved by tailoring the process parameters and thus process optimization studies are still being carried out for solid parts from different alloys [17]. Triply periodic minimal surfaces (TPMS), honeycomb are some of the examples of porous metal structures that are desirable for applications in aerospace, automobile, heat exchangers but their fabrication strategies are not well established [18]. The porous structures consist of multiple pores organized in unit cell configurations throughout the volume of the specimen.

The porous structures consist of multiple pores organized in unit cell configurations throughout the volume of the specimen. The design of the TPMS structures is slightly complicated as they are mathematically defined surfaces, while the lattices and the honeycomb structures can be easily designed with the help of various CAD packages [19].Gibson et al., developed a model known as the Gibson-Ashby model to determine the relative density based on the lattice structure’s theoretical mechanical properties [20]. Zhang et al. used SLM technique to fabricate 316L TPMS structures and studied their mechanical and energy absorption capabilities [21]. Abueidda et al., studied the effect of porosity levels on the mechanical properties of Selectively laser sintering (SLS) fabricated specimens [22]. In a similar study, Yang et al., compared the experimental values of elastic modulus and yield strength of SLM fabricated Ti64Al4V samples with that of finite element simulations [23].

In a similar study, Yang et al., compared the experimental values of elastic modulus and yield strength of SLM fabricated Ti64Al4V samples with that of finite element simulations [23]. Numerous studies have been conducted to examine the correlation between the porosity grading methodologies and the physical properties [24-26]. Functionally graded diamond-type lattice structures and their mechanical properties were studied by Han et al [24]. It was observed that the diamond structured lattice failed layer by layer as opposed to the diagonal shear band that is commonly observed in compressively loaded homogenous lattices [24] . Similar results were observed on the functionally graded 316L gyroid lattice [27]. Wang et al., investigated the effect of varying cell sizes on the mechanical properties of SLM fabricated Ti6Al4V alloy [28]. It was found that the stress-strain responses revealed a post-peak significant drop in the strength.

It was found that the stress-strain responses revealed a post-peak significant drop in the strength. To our knowledge, no comprehensive study has been done to compare different functionally graded porous metal lattice fabricated by 3D printing. In this work, different functionally graded TPMS structures are designed from 316L using SLM Solutions 125 HL printer. Compressive tests were performed on the as-built 316L lattices to determine the mechanical properties of the TPMS structures. Part densities of the samples were measured and their effect on the mechanical properties was evaluated.2. MATERIALS AND METHODS2.1 Design of Lattice StructuresA total of 5 TPMS structures (i.e., Diamond, Gyroid, Primitive, Neovius, and Fisher-Koch S) were designed with the help of MS Lattice, an open source TPMS surface generator [29] as shown in Figure 1. The total size of the samples was maintained at 8 mm x 8 mm x 8 mm with each unit cell being 2 mm x 2 mm x 2 mm.

The total size of the samples was maintained at 8 mm x 8 mm x 8 mm with each unit cell being 2 mm x 2 mm x 2 mm. The porosity of the samples was varied from 30% to 70%. The TPMS were then converted into 3D stereo-lithography (STL) files. The STL files were imported into Solidworks (Dassault Systèmes, Vélizy-Villacoublay, France) and two plates (each with the thickness of 1 mm) were added at the top and bottom of the porous structures to support the mechanical testing. The Solidworks files were imported to Materialise Magics (© Copyright Materialise 2021, Leuven, Belgium) to assign the supports and corresponding process parameters2140for printing.

Designed TPMS structures with varying porosity levels from 30% to 70% and build direction (BD) indicated by the arrow and types: a) Diamond; b) Gyroid; c) Primitive; d) Neovius; e) Fisher-Koch S.2.2 Powder Preparation and FabricationThe 316L TPMS specimens were fabricated using SLM 125 HL LPBF printer (SLM Solutions Group AG, Lübeck, Germany). The SLM 125 HL LPBF printer is equipped with a 400 W Ytterbium fiber laser and has a build volume of 125mm×125mm×125mm. Gas atomized 316L powder was obtained from SLM solutions AG. Figure 2 shows the scanning electron microscope2141(SEM) micrograph of the fresh 316L powder, and its composition is listed in Table 1. From the SEM micrograph it can be seen that the powder particles are mostly spherical.Figure 2.

From the SEM micrograph it can be seen that the powder particles are mostly spherical.Figure 2. a) SEM micrograph of fresh 316L powder used during fabrication and b) Schematic depicting the stripes scan strategy used during the SLM process.Table 1: The chemical composition of 316L powder obtained from SLM Solutions.ElementFeCrNi Mo Nb + Ta Mn SiPSCN OMass fraction (%)Balance 16-18 10-142-3210.045 0.03 0.03 0.1The samples were fabricated with the SLM Solutions suggested process parameter set. The laser power (LP = 200 W), scan speed (SS = 800 mm/s), hatch spacing (HS = 120 μm), layer thickness (LT = 30 μm) and the stripes scan strategy as displayed in figure 2 were employed for the fabrication of TPMS samples.

The energy density (Ev) used during the fabrication process is: 69.45 J/mm3 calculated using equation [9, 11, 15, 16]:Equation 6𝐿𝐿𝐿𝐿 𝑆𝑆𝑆𝑆 × 𝐻𝐻𝑆𝑆 × 𝐿𝐿𝐿𝐿𝐸𝐸𝑣𝑣 =2.3 Experimental ProcedureA wire electrical discharge machine was used to remove the as-built samples from the stainless-steel substrate. Density measurements were performed using the Archimedes method. The samples were first weighed in air with the help of a Precision Balance. The samples were later measured in water and the density was calculated. The quasi-static compression tests were performed at room temperature using an Instron 6559 universal testing machine equipped with a 50 kN load cell. A constant strain rate of 10-2 sec-1 was applied perpendicular to the build direction. The maximum displacement allowed for each sample was limited to 50% or 5 mm due to use of the 50 kN load cell.3.

The maximum displacement allowed for each sample was limited to 50% or 5 mm due to use of the 50 kN load cell.3. Results and Discussion3.1 Part Density Analysis2142The mechanical properties of the TPMS specimens depend significantly on the density of the fabricated surface. Figure 3 shows the part density values for the fabricated TPMS specimens. The highest part density values were found for the Fisher-Koch S (98.77%) lattice followed by the Diamond (97.85%) and Gyroid (97.69%) lattices. The Neovius and the primitive lattice structures yielded comparatively lower part density values of 96.28% and 93.31% respectively. The achievement of near fully dense samples is usually expected from the LPBF process as they are capable of fabricating samples with complex geometries. The samples were fabricated with the optimized processing parameters for 316L as suggested by the SLM solutions, given in Section 2.2.

The samples were fabricated with the optimized processing parameters for 316L as suggested by the SLM solutions, given in Section 2.2. The functionally graded porosities seem to have an impact on the density, and it must be noted that the process parameters and scan strategy play a significant role in determining the density of the fabricated specimens [30]. Therefore, it is important to establish the effect of these processing parameters on the part density of functionally graded lattice in future research. Due to the functionally graded nature of the fabricated specimens and the small unit cell size of 2 mm, it was noticed that the part density values were significantly lower and reduced the manufacturability of these samples as reported by Rajgopalan et al. [31]. The lower part density values can also be due to shrinkage, lack of fusion between powder particles and other defects such as unmelted powder particles.

[31]. The lower part density values can also be due to shrinkage, lack of fusion between powder particles and other defects such as unmelted powder particles. This can be observed at the more porous regions of the sample (70% porous side) as it had bigger overhang sections and deformation due to thermal stresses [32, 33].Figure 3. Part densities of the 316L TPMS specimens with same unit cell size of 2mm.3.2 Quasi-Static Compression Testing The part density and unit cell type of the functionally graded lattice specimens govern the mechanical properties. The stress-strain curves of as-built TPMS samples are shown in Figure 4. The tests were performed with the loading direction perpendicular to the build direction ( See Figure 1). The Young’s modulus was obtained by calculating the slope of the elastic region and the compressive yield strength of samples was determined by using 0.2% offset approach [32].

The Young’s modulus was obtained by calculating the slope of the elastic region and the compressive yield strength of samples was determined by using 0.2% offset approach [32]. The stress-strain curves indicate that after the initial elastic region, the structures deform plastically and continue to absorb energy. This is followed by the densification stage, where the samples behave similar to that of solid material as there is a large self-contact area as seen in Figure 6. A similar trend was observed in a study conducted by Li et al. [33]. From Figure 5, we can see that the Fisher-Koch S specimen had the highest Young’s modulus value of 5.12 GPa followed by the2143Diamond (4.51 GPa) and Gyroid (4.21 GPa) lattice samples. The Primitive and Neovius structues resulted in lower Young’s modulus values of 2.48 GPa and 2.13 GPa respectively. However, the compressive yield strength was highest for the primitive structure (112.83 MPa) followed by the Fisher-Koch S (98.66 MPa) sample.

However, the compressive yield strength was highest for the primitive structure (112.83 MPa) followed by the Fisher-Koch S (98.66 MPa) sample. The diamond and gyroid samples had slightly lower yield strength values of 80.41 MPa and 69.47 MPa. The neovius structure resulted in the lowest yield strength value of 42.52 MPa. The experimental deformation of the samples at different strains is presented in Figure 6. The samples did not deform with a diagonal shear but rather deformed layer by layer as in the case for uniformly graded lattices [24, 27].Figure 4. Stress-Strain curves of the different as-fabricated functionally graded lattice specimens.Figure 5. Young’s Modulus and yield strength values of the different as-fabricated functionally graded lattice specimens.2144Figure 6. Experimental deformation of a) Diamond, b) Gyroid, c) Neovius, d) Primitive, and e) Fisher-Koch S at different strains.4.

Experimental deformation of a) Diamond, b) Gyroid, c) Neovius, d) Primitive, and e) Fisher-Koch S at different strains.4. CONCLUSION In this study, functionally graded lattice 316L structures with porosity varying from 30% to 70% were designed and fabricated using the LPBF technique. The fabricated lattice specimens were evaluated on part density and quasi-static compression behavior. The part density of the Fisher- Koch S, Diamond and Gyroid lattice samples were closer to the desired part density with no internal voids i.e., 100% dense. The Primitive lattice sample had the lowest part density value. The primitive structure resulted in a higher compressive yield strength value of 112.834 MPa. On the contrary, it had a low Young’s modulus of 2.485 GPa. The fisher-Koch S sample resulted in a high2145value for Young’s modulus (5.1255 GPa) and had the second highest yield strength (98.66 MPa) value.5.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEVALUATION OF LIQUID DOPING METHODS FOR USE IN LASER POWDER BED FUSIONT. M. Davis* and N. B. Crane*Department of Mechanical Engineering, Brigham Young University, Provo, UT 84602AbstractLaser powder bed fusion (LPBF) is an additive manufacturing (AM) process that is well known for its geometric versatility and high-quality parts. While the properties of LPBF parts are commonly superior to those made using other AM techniques, LPBF is generally limited to a single material in any given build. While LPBF can accommodate the integration of multiple components into a single part geometrically, the material limitation leads to over-designing to ensure that every component can complete their various functions.

Some studies have shown potential methods of 3D composition control throughout a part, but these methods are subject to high cost increases due to build time increases and decreased powder recyclability. A new approach to multi-material LPBF uses liquid dopants to alter the composition in location-specific areas. The current study evaluates two different liquid deposition methods – direct write and inkjet deposition – in relation to their adaptability and utility in LPBF. Inkjet deposition is shown to have significant benefits compared to the direct write method.IntroductionAdditive manufacturing (AM) – and laser powder bed fusion (LPBF) in particular – is well known for its ability to construct geometric forms that would not be possible using standard manufacturing techniques [1-3].

This geometric versatility has inspired a design practice known as functionality integration – where multiple components are condensed to create a single part that performs all of the functions that were previously attributed to the individual components. While this practice can be quite useful in terms of weight reduction, part count reduction, and meeting spatial confinements; there are also serious drawbacks [4]. One of these drawbacks occurs in LPBF when different functions integrated into a single part have extremely different requirements (e.g. high fracture toughness on the interior and high hardness on the exterior), but the process can only sustain a single material [5]. This dilemma generally results in over-designing some areas of the part to compensate for the compromise in material choice. Over-designing can lead to decreased functional efficiency, increased weight, decreased fatigue life, etc. in LPBF parts.

Over-designing can lead to decreased functional efficiency, increased weight, decreased fatigue life, etc. in LPBF parts. Creating methods to control the material composition spatially throughout a build would allow for designers to mitigate the negative effects and experience the full benefits of functionality integration.The literature reports multiple attempts to expand LPBF processes into the multi-material regime. One approach is to simply switch out the powder feedstock for a different material at a certain point in the build [6, 7]. While this has been done successfully in some cases, it greatly increases the overall build time. Because of the time sacrifice, it is generally limited to a single, large composition change during the build. This discrete change in composition can be problematic.

Because of the time sacrifice, it is generally limited to a single, large composition change during the build. This discrete change in composition can be problematic. Materials with vastly different properties create extra residual stresses from thermal expansion and contraction during fabrication and poor adhesion at the inter-material surface due to poor wetting and insufficient mixing [7, 8]. Some studies have shown that adhesion can be843improved by remelting the inter-material zone multiple times to improve mixing between the dissimilar materials, but this adds even more time to the process [7]. In addition to quality concerns, this method is still limited to material change in one direction and few changes throughout the part.Other research has been done on methods of three-dimensional variation of composition. Wei et al.

Wei et al. [9] have implemented a process that uses a vacuum to remove base powder in select areas and a separate hopper and nozzle to deposit powders of different compositions in the excavated areas. This novel method has been shown to work with glass-metal, metal-metal, ceramic-metal, and metal-polymer systems and has shown significant improvement in terms of spatial control throughout all three dimensions [9-14]. The method of Wei et al. has the downside of allowing powder of different compositions that is not fused to remain in the powder bed and contaminate the unfused powder that would be recycled. Since powder recycling is one of the main factors that contributes to the economic feasibility of LPBF, this method significantly increases the cost of the overall process [15]. The other significant cost issue that is affected by this multi- material method is the increase in time necessary to complete each layer.

The other significant cost issue that is affected by this multi- material method is the increase in time necessary to complete each layer. Vacuuming out and replacing powder can more than double the overall process time that originally consisted of only powder spreading and laser scanning. Large composition changes in three dimensions present the same issues observed for large composition changes in one dimension.New Approach to Multi-material Laser Powder Bed FusionMany of the issues that are present in these current multi-material LPBF (MM-LPBF) efforts could be mitigated or eliminated by using liquid or liquid-encased dopants as the means of altering composition in a way that does not significantly add to the build time. The nature of the LPBF process provides two distinct opportunities for the introduction of liquid dopant to the powder bed. The first of these opportunities occurs after a layer is fused by the laser but before the next layer of powder is spread.

The first of these opportunities occurs after a layer is fused by the laser but before the next layer of powder is spread. The second is just after the powder is spread but before the layer is fused.In the first case, liquid dopant could be deposited on the solid substrate using either a direct write (controlled extrusion through an orifice) or droplet-based deposition method like inkjet printing (see Figure 1). The liquid would be dried either by the heat in the build chamber or using an exterior heat source (such as a lamp or the laser at low power) after which the next layer of powder would be spread. This method provides a distinct advantage of the dopant being covered by powder of the base material, which would reduce the amount of dopant that is evaporated or ejected as spatter during the fusing process.

Since the dopant is initially concentrated in a specific area (between the powder and substrate) this may also affect the final location and dispersion of the dopant throughout the melt pool. Direct write can also be used with a wide range of suspension properties and has little risk of clogging in the nozzle. However, direct write can be slow since it is generally only used with a single nozzle.844Figure 1 - Direct write liquid dopant deposition method. The liquid dopant is deposited using an open-tip needle directly onto the previous (fused) layer. Once the dopant is dry, the powder is spread over the top and the fusing process is carried out and the base and dopant materials are mixed.In the second case, the powder layer would be spread and then the liquid dopant would be deposited into the spread powder before the layer is fused (see Figure 2).

Deposition in this method could only be done using a droplet-based system like the inkjet printing technology used in binder jetting – another AM technique. An inkjet printhead could be mounted on the coater blade system to print while spreading the next layer of powder – minimizing the impact on the total build time.Figure 2 - Inkjet liquid dopant deposition method. After the powder is spread, liquid dopant is deposited into the powder in the form of micro-scale droplets from an inkjet printhead. Once the dopant and powder bed are dry, the fusing process is carried out.This work focuses on evaluating the feasibility of depositing liquid dopants as part of the LPBF process by exploring the implications of exploiting the two deposition options described. These situations are replicated in a way that does not permanently alter the existing LPBF system.

These situations are replicated in a way that does not permanently alter the existing LPBF system. Direct write of a liquid dopant onto a solid substrate mimics the deposition onto a fused layer before the next layer of powder is spread, while inkjet deposition into a powder bed represents deposition after a powder layer is spread but before the layer is fused together. Simplified versions of these two methods are performed with careful observation of any requirements or outcomes that could significantly affect integration of the method to the LPBF process. Conclusions are made as to the deposition quality, integration feasibility, and system productivity of the two methods in relation to their application in LPBF.845MethodsMaterialsTwo dopant material systems were selected based on the availability of stable commercial suspensions and the potential for property enhancement in SS 316L.

One material system that has previously been studied with LPBF is alumina (Al2O3) reinforcement in a stainless steel 316L matrix [16]. In this study, Li et al. showed that the addition of 1 to 3 𝑤𝑡% alumina to stainless steel processed using LPBF improved hardness in all cases and improved yield and tensile strengths in the case of 1-𝑤𝑡% alumina addition. While this study by Li et al., and most multi-material studies in LPBF, have been done using physically mixed matrix and additive powders, the property improvements shown demonstrate that the alumina/SS 316L material system is a useful material system for testing composition adjustment using alternative doping techniques.Another material system of interest is zirconia (ZrO2) reinforcement in a steel matrix. Koopmann et al.

Koopmann et al. [7] showed that zirconia powder can be processed reasonably well using LPBF and that it can have good adhesion with steel when the interface between the two materials undergoes sufficient mixing during the laser processing. While their research was focused more on large composition changes, the good mixing between the two materials and processability of zirconia indicate that it could be a suitable test subject for feasibility of composition adjustment using alternative doping techniques.In both the direct write and inkjet deposition methods, SS 316L powder (CL 20ES, Concept Laser, 29.9 𝜇𝑚, spherical) is used as the base material, to which either alumina or zirconia is added in the form of a water-based slurry using one of the two deposition methods described. The alumina slurry is Gamma B 0.05 𝜇𝑚 Alumina from LECO with a 10 𝑤𝑡% concentration alumina with an added 10 𝑤𝑡% propylene glycol.

The alumina slurry is Gamma B 0.05 𝜇𝑚 Alumina from LECO with a 10 𝑤𝑡% concentration alumina with an added 10 𝑤𝑡% propylene glycol. The zirconia slurry is ZR100/20 from NYACOL with 20 𝑤𝑡% colloidal zirconia with a mean particle size of 100 𝑛𝑚. The alumina and zirconia concentrations were measured to be 10.6 and 25.4 𝑤𝑡% respectively. The LPBF machine used is a Concept Laser M2 Cusing Multilaser.Direct WriteDirect write deposition describes a system in which a material is extruded directly onto a substrate through an orifice (see Figure 1 and Figure 3). The material creates a liquid bridge with the substrate that moves with the needle and leaves a liquid track behind. This technique would be difficult to apply to a powder bed as the fluid meniscus would likely move the powder during liquid deposition; however, it is suitable for deposition on a solid substrate.

This technique would be difficult to apply to a powder bed as the fluid meniscus would likely move the powder during liquid deposition; however, it is suitable for deposition on a solid substrate. In the current study, the direct write deposition is applied to deposit alumina or zirconia slurry onto a solid plate made of SS 316L to simulate the case in which dopant is deposited on a previously fused layer.A custom direct write system (see Figure 3) was used for printing [17-19]. The system uses a 3-axis CNC stage and a stepper motor to control the plunger of a syringe with an open-tip needle to deposit the slurry. The needle used was 25 gauge with a 0.305 𝑚𝑚 inner diameter. The plate was leveled with respect to the x- and y-axis movement of the system and the needle was zeroed to the surface of the flat plate. Deposition was performed with the needle tip at a distance of 0.10 to 0.15 𝑚𝑚 above the plate surface.

Deposition was performed with the needle tip at a distance of 0.10 to 0.15 𝑚𝑚 above the plate surface. Areas and lines were deposited at a flow rate of 0.0454846𝑚𝑚3/𝑚𝑚 while the needle traveled at speeds of 80 𝑚𝑚/𝑠 relative to the plate surface. Area depositions were performed in a concentric, rectangular pattern. The spacing between tracks was varied to control the total amount of slurry deposited in a given area. The dopant concentration in the slurries was also varied by adding distilled water to dilute the slurry to a predetermined concentration. The combination of these two adjustments was used to vary the overall dopant concentration amounts as described in Table 1 for the alumina depositions and Table 2 for the zirconia depositions.Figure 3 - Direct write deposition system (image from Romero et al. [19] used with permission).

Using the LPBF machine’s coater system and build chamber controls, a847layer of powder was spread over the dried alumina by adding 500 𝜇𝑚 of powder and removing powder in 100 and 50 𝜇𝑚 increments to get to a 50 𝜇𝑚 powder layer. This was done to determine if the alumina deposits were stable during the powder coating step or if they would break down and contaminate the powder.Direct write single-line depositions were also performed to investigate the difference between small- and large-area depositions. In these cases, the substrate was heated during deposition to a temperature of 50 °𝐶 rather than baking the substrate after deposition to simulate a heated build plate during the LPBF process.InkjetThe inkjet method is suitable for printing dopants into spread powder layers. While printing into powder can be challenging [20, 21], it is successfully done in both the binder jetting and multi jet fusion/high speed sintering commercial processes.

While printing into powder can be challenging [20, 21], it is successfully done in both the binder jetting and multi jet fusion/high speed sintering commercial processes. In order to simulate dopant deposition onto loose powder, a solid SS 316L plate was used as a base onto which walls were built to create isolated pockets of powder. These pockets reduce the error in quantifying the amount of dopant deposited by limiting the area into which the dopant can spread. The pocket walls were built in the LPBF machine with a laser power of 370 𝑊, a scan speed of 1350 𝑚𝑚/𝑠, and a spot size of 130 𝜇𝑚. The walls were built by depositing a 25 𝜇𝑚 layer of powder, fusing the pocket walls, depositing another 25 𝜇𝑚 powder layer (50 𝜇𝑚 total) and fusing the pocket walls again (see Figure 4). These 25 𝜇𝑚 increments are smaller than typical to ensure total fusing of the pocket walls with the base plate so that each pocket was totally separate from the others.

These 25 𝜇𝑚 increments are smaller than typical to ensure total fusing of the pocket walls with the base plate so that each pocket was totally separate from the others. Once these pockets were built, the powder inside the pockets was left undisturbed while the plate was taken out of the LPBF machine and transported to the inkjet printing station.Figure 4 - Pockets walls printed 50 𝜇𝑚 tall on substrate to separate powder beds of the same height. Powder beds before inkjet deposition are fairly uniform. Pocket walls and individual powder beds are labeled. Individual pocket labels are also included for future reference.A simple inkjet setup (see Figure 5) was used to deposit colloidal zirconia slurry into the walled-off powder beds [20, 21].

Individual pocket labels are also included for future reference.A simple inkjet setup (see Figure 5) was used to deposit colloidal zirconia slurry into the walled-off powder beds [20, 21]. This setup uses a pressurized chamber connected to a single, 80848𝜇𝑚-nozzle, piezo-electric, drop-on-demand print head (MicroFab Technologies, Inc.; Part # MJ- AB-01-80-8MX) controlled in coordination with the movements of the stages to create lines of consistently-spaced droplets. The nozzle released droplets at a rate of 1000 𝐻𝑧 while moving in the x-direction at a speed of 60 𝑚𝑚/𝑠 creating a droplet spacing of 60 𝜇𝑚. The plate was baked after deposition at 180 °𝐶 for 30 minutes to evaporate the liquid from the slurry. When multiple passes were necessary to achieve the desired concentration, the plate was baked between passes as well.Figure 5 - Single-nozzle inkjet deposition system.

When multiple passes were necessary to achieve the desired concentration, the plate was baked between passes as well.Figure 5 - Single-nozzle inkjet deposition system. The three axis stage controls the deposition location while the piezo-electric nozzle controls the droplet size and ejection frequency into the powder bed or onto the substrate.Inkjet CalculationsThe concentration of dopant in the powder bed was controlled by varying the line spacing and number of lines for each pocket (see Table 3). These parameters were chosen based on an estimate of how much dopant would be integrated into the melt pool during laser processing.

These parameters were chosen based on an estimate of how much dopant would be integrated into the melt pool during laser processing. Some of the values are purposefully higher than might be typically needed in order to test the upper limits of the deposition method.Table 3 - Deposition parameters for inkjet deposition of zirconia slurry.Pocket # 1 2 3 4 5 6# of passes 3 3 1 1 1 1# of lines / pass Line spacing (𝒎𝒎) 1060 774 1506 1506 886 2900.0226 0.0310 0.0159 0.0159 0.0271 0.0830849Droplet volume was measured by jetting the zirconia slurry in a stationary position for 10 minutes into a 5 𝑚𝐿 beaker of known mass. The beaker was then baked at 180 °𝐶 for 30 minutes – leaving only the zirconia – and weighed again. Subtracting the original mass of the beaker from the mass of the beaker and zirconia gives the mass of zirconia deposited over the time interval (𝑚𝑧).

Subtracting the original mass of the beaker from the mass of the beaker and zirconia gives the mass of zirconia deposited over the time interval (𝑚𝑧). This can then be used to estimate the amount of slurry deposited (𝑚𝑠𝑙) by 𝑚𝑠𝑙 = 𝑚𝑧/𝑐 where 𝑐 is the zirconia concentration in the slurry. The number of droplets can be calculated as 𝑛𝑑 = 𝑡 ∗ 𝑓 where 𝑡 is the deposition time and 𝑓 is the droplet frequency. The slurry mass and number of droplets can then be used to calculate the mass and volume of an average droplet (𝑚𝑑 and 𝑉𝑑 respectively) by 𝑚𝑑 = 𝑚𝑠𝑙/𝑛𝑑 and 𝑉𝑑 = 𝑚𝑑/𝜌𝑠𝑙 where 𝜌𝑠𝑙 is the density of the slurry provided by the manufacturer (1.22 𝑔/𝑐𝑚3). The average amount of zirconia per droplet (𝑚𝑧𝑑) can also be calculated as 𝑚𝑧𝑑 = 𝑚𝑧/𝑛𝑑 using the same values for 𝑚𝑧 and 𝑛𝑑 as above.

The average amount of zirconia per droplet (𝑚𝑧𝑑) can also be calculated as 𝑚𝑧𝑑 = 𝑚𝑧/𝑛𝑑 using the same values for 𝑚𝑧 and 𝑛𝑑 as above. Using these calculations, the droplets were measured to have a volume of about 0.145 𝑛𝐿/𝑑𝑟𝑜𝑝𝑙𝑒𝑡 and to contain about 35.4 𝑛𝑔 of zirconia in each droplet.The droplet volume described above was used to calculate estimates of the amount of slurry/dopant deposited into the powder beds (see Table 4). Total volume was obtained by calculating the number of droplets in each line and multiplying by the number of lines for the deposition. The total saturation is the measure of the amount of space between powder particles that is filled by the slurry. The powder packing fraction was assumed to be 50%, which is typical for non-compacted powder beds [22]. An area density is also calculated by dividing the total mass of zirconia deposited over the area that it was deposited.

An area density is also calculated by dividing the total mass of zirconia deposited over the area that it was deposited. This total mass of zirconia is obtained in similar fashion to the total volume by using the total number of droplets in a given deposition.Table 4 - Predicted volume, saturation, and density calculations for the inkjet parameters mentioned in Table 3. Pocket numbers are defined in Figure 4.Pocket # Total volume1 2 3 4 5 6deposited (𝒎𝑳) 0.1794 0.1309 0.0850 0.0850 0.0499 0.0163Powder bed saturation (%) 1281.8 935.4 607.2 607.2 356.8 116.5Area density of zirconia deposited (𝝁𝒈/𝒎𝒎𝟐) 78.19 57.06 37.04 37.04 21.76 7.11Results and DiscussionThe resulting deposits from the two deposition methods can be seen in Figure 6 and Figure 7. Since there are no quantitative metrics readily available to evaluate these deposition methods, qualitative visual observations of these resulting deposits are the basis for many of the claims as to the value of the two methods.

Other observations were also made by the authors during the deposition processes that are included in the ensuing discussion.850Figure 6 - Dried alumina (left) and zirconia (right) deposits performed using direct write deposition.Figure 7 - Resulting deposits from inkjet deposition method into small, enclosed powder bed using various dopant densities.Three criteria were selected to evaluate and compare the two deposition methods. These criteria are deposition quality, integration feasibility, and system productivity. Deposition quality refers to the uniformity of the deposited material as well as the accuracy of quantity and location of the deposited material. Irregularities, non-uniformities, and inaccuracies would diminish the value that a deposition method could provide. Integration feasibility refers to how well the deposition method could be integrated into the LPBF process. This includes reliability and repeatability of the deposition method.

Integration feasibility refers to how well the deposition method could be integrated into the LPBF process. This includes reliability and repeatability of the deposition method. It also involves any ways in which the deposition method would disrupt, interfere with, or otherwise negatively impact the LPBF process as a whole. System productivity refers to the overall value that a combined deposition/LPBF system would provide.851This includes build time increases, feasible concentration ranges, and other things that would need to be considered when implementing the full deposition/LPBF system.

The direct write and inkjet deposition methods are both evaluated according to these three criteria by observing various characteristics of the deposition processes in action as well as qualities of the resulting dopant deposits.Deposition QualityDirect WriteThe quality of the direct write process was mostly impacted by inconsistencies observedduring deposition and in the resulting dopant structures after drying.The first non-uniformity was observed in moments when the needle tip was too far from the deposition surface. In these instances, the slurry would “skip” small sections of the deposition area because the slurry would build up around the edge of the needle instead of wetting and depositing on the substrate. The skipping would eventually end when the size of the slurry bead on the needle was large enough to make contact with slurry that had already been deposited or the substrate itself.

The skipping would eventually end when the size of the slurry bead on the needle was large enough to make contact with slurry that had already been deposited or the substrate itself. Areas where skipping occurred can easily be seen in recently-deposited areas as small holes or lines in the deposition areas (see Figure 8 - Photos of liquid zirconia slurry in different concentrations shortly after deposition by direct write. Some instances of “skipping” are indicated by red arrows. A large area in the bottom right deposition did not deposit due to silicone contamination on the surface that altered the plate wetting. ).Figure 8 - Photos of liquid zirconia slurry in different concentrations shortly after deposition by direct write. Some instances of “skipping” are indicated by red arrows.

).Figure 8 - Photos of liquid zirconia slurry in different concentrations shortly after deposition by direct write. Some instances of “skipping” are indicated by red arrows. A large area in the bottom right deposition did not deposit due to silicone contamination on the surface that altered the plate wetting.Skipping was also observed in single-line depositions (see Figure 9). The effect was slightly different, however, as there were generally no nearby depositions that could be used to reestablish contact between the slurry in the needle and the substrate surface. When this happened, much more slurry built up on the needle tip than during the area depositions. This resulted in larger sections of the deposition area without dopant and large bead-like areas shortly after the blank sections.852Figure 9 - Direct write deposition of zirconia in single-track lines. Some instances of “skipping” are indicated by red brackets.

Some instances of “skipping” are indicated by red brackets. Beads resulting from the skipping are indicated with red arrows.Another type of non-uniformity was observed when the dopant material moved toward the center of the deposition areas during the drying/baking stage. As the edges of the deposition areas began to dry, the solid particles were transported inward ahead of the solid/liquid interface [23]. The literature attributes this type of capillary push to the decreasing height of the liquid wedge near the contact line [24, 25]. The result was a higher-concentration area in the center of the deposition area with low dopant concentrations near the edges (see Figure 10). While this trend was consistent for all of the direct write depositions, the magnitude of the resulting dopant concentration disparity varied with the input slurry concentration.

While this trend was consistent for all of the direct write depositions, the magnitude of the resulting dopant concentration disparity varied with the input slurry concentration. When the slurries were diluted to lower concentrations, less of a gradient was observed throughout the deposition area.Figure 10 – Recently-deposited liquid alumina slurry diluted to different concentrations. Each image shows the deposition of a unique concentration to a different quadrant on the substrate. Left to right: 3 𝑤𝑡%, 2 𝑤𝑡%, 1 𝑤𝑡%, and 0.5 𝑤𝑡% slurry. Dopant agglomeration near the centers of the deposition areas is observed as more white (alumina) than the exterior regions of the depositions.One last concern about the quality of the direct write method is the structure and stability of the resulting dopant deposit – especially in the highly-concentrated regions. In the alumina deposits, the highly-concentrated centers were quite stable despite some small cracks.

In the alumina deposits, the highly-concentrated centers were quite stable despite some small cracks. The zirconia deposits, however, formed extremely fragile, highly discontinuous structures in their centers that delaminated from the substrate in most cases (see Figure 6). These structures would be likely to break and move under almost any application of force. Since this structure fragility was only seen in the zirconia direct write depositions, this can be categorized as a reflection on the properties of the slurry/dopant and the drying conditions.Overall, the deposition quality of the direct write method is not promising as it is non-uniform and can produce structurally unsound deposits in the case of zirconia dopant.853InkjetWhile the inkjet deposition method has the potential to have great uniformity, there are also many opportunities for error.

In general, the samples with dopant deposited using this method are fairly uniform (see pockets 3, 4, and 5 in Figure 11) as long as simple errors can be avoided. One of these errors is clearly seen near the bottom of pocket 4 and just above the center of pocket 3. In these cases, the nozzle was clogged temporarily during the deposition resulting in a few missed lines where dopant should have been deposited. Another error seen in pocket 5 is contamination from an outside source as flakes of residue material fell from the nozzle structure into the pocket during deposition. These errors all affect the uniformity of the resulting deposition; however, they could be resolved with a more controlled inkjet system commonly used in industrial applications.Figure 11 - Zirconia-doped powder beds at different concentrations with pocket 1 being the highest concentration and pocket 6 being the lowest (see Table 4 for detailed concentration estimates).

Contaminates and skipped lines are the result of user-error during deposition while increased roughness and non-uniformity of dopant concentration are important results that will impact the LPBF process.While the mid-range depositions appear uniform, the higher-density depositions (see pockets 1 and 2 in Figure 11) show more variance in the uniformity of the dopant concentration throughout the pocket. This is due to the increased movement of the liquid and powder during the deposition stage.

This is due to the increased movement of the liquid and powder during the deposition stage. The greater amount of liquid in the pocket created a pool of liquid, in which the 316L powder particles could flow.The uniformity of the inkjet method can be somewhat limited in the higher range of saturations; however, increasing the dopant concentration in the slurry or depositing the material in multiple passes with drying between passes could expand the feasible range of doping deposits.Integration FeasibilityDirect WriteIn practice, there are a few things that would need to be considered before implementingthe direct write method in an LPBF system.854The first of these is the needle tip distance from the surface. During the direct write operation, it was observed that the skipping mentioned above could be avoided by keeping the tip of the needle within about 0.15 𝑚𝑚 of the substrate surface during deposition.

During the direct write operation, it was observed that the skipping mentioned above could be avoided by keeping the tip of the needle within about 0.15 𝑚𝑚 of the substrate surface during deposition. This requires an extremely flat substrate which is not always the case with built layers in the LPBF process. During the depositions, it was noted that little to no skipping occurred when the substrates were level to within about 0.05 𝑚𝑚.The incorporation of a direct write system into an LPBF machine would also require significant modifications. This would entail an additional motion axis for the nozzle to move in the direction perpendicular to the coater blade or two motion axes if the system moves independent from the coater system.

This would entail an additional motion axis for the nozzle to move in the direction perpendicular to the coater blade or two motion axes if the system moves independent from the coater system. A vertical motion axis may also be warranted to ensure that the nozzle tip does not come in contact with the metal powder at any point in the process and that it can be lowered to the correct height as mentioned previously.Another important consideration is whether the dopant will contaminate the powder that will not be fused as part of the final build. The delaminated flakes of dried zirconia slurry would be broken off during powder spreading and could travel to almost any part of the build chamber. This would be problematic as the composition could be changed in the wrong locations within the part, unfused powder could be contaminated, and powder flowability could decrease due to the different shape and size of the zirconia inclusions.

However, this problem is specific to the properties of the dopant. The plate with alumina depositions was inserted into the LPBF machine and powder was spread over the top. After removing the plate and clearing the powder, it was determined that the powder spreading process did not cause any visible physical damage to the deposited alumina.Powder contamination could also be a concern when the dopant is still in liquid form. If the liquid in the slurry does not evaporate before the next layer of powder is spread, the powder could be contaminated as mentioned above. The moisture could also severely limit the flowability of the powder or even cause build-ups on the coater blade.

The moisture could also severely limit the flowability of the powder or even cause build-ups on the coater blade. Both of these cases could initiate the formation of defects that would propagate through multiple layers if not the entire build [26-28].InkjetThere are a few things that were noted in the inkjet process that may affect the way it isimplemented in LPBF applications.One of these observations was the dramatic change in color in the higher-density depositions. While all of the depositions did show some change in color, the change in pockets 1 and 2 were the most pronounced. This may be due to the dopant being present on the top surface of the powder bed at higher concentrations compared to the other depositions. This is of particular importance for LPBF because the added material being on top of the material could change the effective absorptivity of the powder bed.

This is of particular importance for LPBF because the added material being on top of the material could change the effective absorptivity of the powder bed. This change would require a shift in processing parameters to achieve the same part quality.Another feature that could impact the LPBF process is the powder bed roughness observed in pocket 6 (see Figure 11). In binder jetting, this increase in roughness has been described as a first layer phenomenon that occurs when the droplets are too widely spaced [29]. This can cause855poor lamination throughout the first few layers in binder jetting. Similarly, in LPBF, increased roughness in one layer can cause balling and other defects in subsequent single-track scans or fused layers and can propagate throughout the part [26-28].

Similarly, in LPBF, increased roughness in one layer can cause balling and other defects in subsequent single-track scans or fused layers and can propagate throughout the part [26-28]. This could prove prohibitive to low- saturation depositions using the inkjet method; however, slurries or solutions with lower dopant particle densities could be used to achieve smaller composition changes while maintaining the same saturation levels.While these effects on the LPBF process may require some adjustment, they are not prohibitive as other processes (such as binder jetting) have successfully addressed them through printing parameter adjustments [20, 21, 29].Inkjet printheads can also be fitted across a wide span and could easily cover the entire width of the coater blade.

This would make it so that no additional motion axis is required for incorporation into the LPBF system.System ProductivityDirect WriteWhile using liquid dopants to alter composition may be quite attractive in an LPBF setting, the benefits of the method being used must outweigh the associated costs to make a reasonable claim at increasing value.The main cost that this method incurs is the extra time that it will take. Since the deposition step cannot occur simultaneous to any of the other LPBF process steps, every layer where dopant is needed would take considerably longer to complete. This time addition could be in the range of a few seconds to a few minutes for each layer depending on the traverse speeds, the amount of dopant to be deposited, and the area over which it was to be deposited. The dopant then also needs to dry before continuing – to avoid introducing moisture to the delicate powder process as mentioned above.

The dopant then also needs to dry before continuing – to avoid introducing moisture to the delicate powder process as mentioned above. This wait could add more seconds or minutes onto each layer time depending on the size and shape of the deposits.The benefits are also limited by the non-uniformity of the resulting dopant deposits. In order to maintain a somewhat uniform deposition, the direct write method as performed in this study would be limited to less than 0.6 𝜇𝑔/𝑚𝑚2 of alumina (or 0.3 𝑤𝑡% of the deposited material when measured with a 50 𝜇𝑚 layer of powder) for any area depositions. While this may be useful to alter some properties, it is on the low side of small composition change. Studies have shown that useful small composition changes are generally between 1 and 5 𝑤𝑡% for dispersion strengthening [16, 30].

Studies have shown that useful small composition changes are generally between 1 and 5 𝑤𝑡% for dispersion strengthening [16, 30]. However, the amount of dopant required to affect other changes in a material may be significantly higher.In this case, the added time costs and small concentration range are extremely limitingwhen considering implementation of a direct write deposition system into a LPBF process.InkjetThe inkjet system has some clear advantages over the direct write system. The added time cost for this method will be close to nothing. Since inkjet printhead arrays can be scaled to perform856at very fast speeds and the printhead can operate simultaneous to the powder spreading, the deposition step will add very little if any time to the overall process.

Considerations may need to be made for drying the liquid dopant; however, a heated build chamber could reduce that time cost significantly as well.Though there are many benefits to using an inkjet deposition system with LPBF, there are some limitations as well. One specific limitation found during this study is that the alumina slurry did not work in the inkjet setup. With the small-aperture piezo-electric nozzles that were used with the inkjet system, a 5.00 𝜇𝑚 filter was used to keep larger particles from damaging the nozzles. This filter made it impossible for the alumina slurry to pass through the system, meanwhile the zirconia slurry had no problems. This illustrates the need for careful preparation of the dopant to be used with an inkjet deposition system.Generally, the inkjet deposition system has the potential to add significant value to the LPBF process.

The problems and limitations can be addressed if the property enhancements are sufficient.ConclusionsOverall, inkjet printing is a better option for liquid-dopant deposition in LPBF because of its superior uniformity, high resolution, minimal time addition to the process, and its history of use with powder beds (e.g. binder jetting). The direct write method could be beneficial in specific circumstances (e.g. single tracks of dopant), but would be much more difficult to implement across the rough surfaces often encountered in LPBF. The only clear advantage that the direct write method could claim over inkjet deposition is that the dopant is underneath the powder and would have a minimal effect on laser absorption. The inkjet method can deposit liquid dopants in concentrations between about 10 and 40 𝜇𝑔/𝑚𝑚2 while the direct write method is limited to less than 0.6 𝜇𝑔/𝑚𝑚2 for large areas.

The inkjet method can deposit liquid dopants in concentrations between about 10 and 40 𝜇𝑔/𝑚𝑚2 while the direct write method is limited to less than 0.6 𝜇𝑔/𝑚𝑚2 for large areas. Choice of material that is suitable for the different methods is of paramount importance as zirconia proved to not be useful in the direct write method while the alumina slurry could not be used in inkjet printing. This limitation may be overcome by using inks that are formatted specifically for inkjet printing.857References[1] A. du Plessis, C. Broeckhoven, I. Yadroitsava, I. Yadroitsev, C.H. Hands, R. Kunju, D. Bhate, Beautiful and Functional: A Review of Biomimetic Design in Additive Manufacturing, Addit Manuf 27 (2019) 408-427. [2] A. Kover, Transformation in 3D: How a Walnut-Sized Part Changed the Way GE Aviation Builds Jet Engines, GE, GE Reports, 2018. [3] J.O.

aluation of Mechanical Properties and Microstructure for Laser Deposition Process and Welding ProcessYaxin Bao1, Jianzhong Ruan1, Todd E. Sparks1, Jambunathan Anand2, Joseph Newkirk2, Frank Liou1 1Mechanical Engineering, University of Missouri-Rolla 2Material Science & Engineering, University of Missouri-Rolla Reviewed, accepted September 14, 2006AbstractLaser Aided Manufacturing Process (LAMP) can be applied to repair steel die/molds whichare currently repaired using traditional welding process in industry. In order to fully understand the advantages of laser deposition repair process over traditional welded-repair process, the mechanical properties such as tensile strength and hardness of H13 tool steel samples produced by these two processes were investigated. The microstructure and fracture surface of the samples were analyzed using optical microscope and SEM (Scanning Electron Microscope).

The microstructure and fracture surface of the samples were analyzed using optical microscope and SEM (Scanning Electron Microscope). Moreover, depositions on substrates with different shapes were studied to evaluate the performance of LAMP on damaged parts with complicated geometric shape.Keywords: Laser deposition, Microstructure, H13 tool steel, RepairIntroductionA laser-based die repair system [1] is being developed in the Laser Aided ManufacturingProcess (LAMP) lab of University of Missouri-Rolla. For die casting, the primary failure mechanism of steel molds is abrasive wear [2]. If a repaired die can provide a similar life to a new one, then it is much more cost-effective to repair it than manufacturing a new die. Currently, the most common used repair process in industry is fusion welding. However, welding has some significant disadvantages, such as the randomicity of manual process which affects the reliability and quality of the repair very much.

However, welding has some significant disadvantages, such as the randomicity of manual process which affects the reliability and quality of the repair very much. On the other hand, the LAMP system automatically complete the repair process with the hybrid laser metal deposition and machining process, and thus, the performance of the repaired part is reliable.This paper concentrates on characterizing laser deposition process compared to the TIG welding. The same substrates were separated into two groups, one group was TIG welded, the other laser deposited. Tensile tests and hardness tests [3] were performed. A fracture analysis and microstructure comparisons were conducted on the samples after mechanical testing. Due to the faster cooling rate, the average hardness of deposited samples is higher than that of welded samples.

Due to the faster cooling rate, the average hardness of deposited samples is higher than that of welded samples. The fracture surface analysis shows that the deposited samples retain some ductility, indicating that the hardness can be improved via heat treatment.280Moreover, in order to test this hybrid system’s ability of repairing on the uneven surface, substrates with different height steps were machined with a wire EDM. After deposited on these substrates, samples were sectioned and polished to check if there is any defect and map out the type and position of defects.MaterialAll the substrate material, the powder for deposition and the filler for TIG welding are H13tool steel, which is a common used material for the die/molds in industry. Table 1 shows the nominal composition of H13 tool steel.Element Composition (%)Table1.

Table 1 shows the nominal composition of H13 tool steel.Element Composition (%)Table1. The main composition of H13 tool steel Mn 0.4C 0.4Cr 5.2V 1Mo 1.5Si 1Ni 0.3 BalanceFeLaser-based Repair System FrameworkThe LAMP system mainly consists of a 5-axis CNC machining center, a 1.0 KW Diode laser, a powder feeder system from Bay State Surface Technologies Inc, and a real–time control system from National Instruments. During the depositing, the substrate is fixed on the fixture of the 5-axis CNC. The nozzle through which the laser and metal powder is transmitted is mounted on a vertical linear axis fixed to z-axis of the CNC. The laser is focused on a small area of the substrate and create a melting pool, and the metal powder is delivered by the powder feeder system into the melting pool to create the deposition.

The laser is focused on a small area of the substrate and create a melting pool, and the metal powder is delivered by the powder feeder system into the melting pool to create the deposition. The x, y and z table positions and velocities are regulated via the CNC machining center controller according to the program generated from the CAD model [4]. This hybrid repair system employs 5-axis positioning system which includes of 3 linear axis and 2 rotating axis. The advantage of it over conventional 3-axis positioning system is that it does not require the support material to build overhang features for 3D parts and allows both the deposition and machining in a single set-up for a part even with intricate or hidden features.Experimental ProcedureThe procedure for the comparison between these two processes is outlined below.1) H13 tool steel was deposited on 18mm cube substrates to 36mm height with the hybrid repair process. The parameters for the process are listed in Table 2.

The parameters for the process are listed in Table 2. The size of the samples was decided by tensile test procedure.281Laser Power700WTable 2. Parameters of Hybrid Repair Process Nozzle Travel Speed 381 mm/minThickness per Deposition Layer 0.5mmPowder Feed Rate 12 g/minHeat Treatmentno2) H13 tool steel was deposited on the same substrates to the same height with the TIGwelding process. This process was completed by Spartan Light Metals Inc. The only difference between procedure used in this project and normal practice is the height of the welding is higher than that used in practice. There is no heat treatment for these samples.3) The samples were cut into tensile “dog bone” shapes with a wire EDM. 4) The tensile test was performed on an Instron 4469 UTM (Universal Testing Machine). 5) The hardness was measured with reference to position within the deposition/welding. 6) The fracture surface was observed with Scanning Electron Microscope (SEM).

5) The hardness was measured with reference to position within the deposition/welding. 6) The fracture surface was observed with Scanning Electron Microscope (SEM). 7) The samples were polished, etched with 2% nital, and then observed for the microstructure with optical microscope.The procedure for the study about substrates with steps is outlined below.1) The substrates were cut with wire EDM. And the steps are 0.01mm, 0.05mm, and 0.1mm. 2) H13 tool steel was deposited on these substrates with the same parameters listed in Table 2. 3) The samples were cut through to show the area of step. 4) The samples were polished, etched with 2% nital, and then observed for the microstructure.Fracture Surface AnalysisFigure 1 is a SEM image of the fracture surface of a specimen which was laser deposited.And figure 2 shows the fracture surface of welding sample.

These images were acquired on a Hitachi S570 SEM under the following conditions: Magnification 40X, Accelerating Voltage 30KV, Working Distance 20mm, Spot Size Con1=Con2=4 [5].From Figure 1, it can be observed that there are both dimpled features and flat features onthe fracture surface of laser deposition sample. Figure 2 shows only flat features for the welding sample. The dimpled features represent ductile fracture, while flat features represent brittle fracture [6]. Therefore, these two images indicate that the laser deposition sample still has some ductility and its hardness can be improved through heat treatment, while the welding sample has little ductility.282dimpled featureFigure 1. Fracture surface of laser deposition sample (300X)Figure 2. Fracture surface of welding sample (300X)283flat featureMicrostructure ComparisonFigure 3. The microstructure sketch of laser deposition sampleFigure 3 shows the microstructure of laser deposition sample.

The microstructure sketch of laser deposition sampleFigure 3 shows the microstructure of laser deposition sample. The microstructure changesgradually from equiaxed structure to dendritic structure over the first four layers and becomes constant with all of the following layers which are showed as one image because all of them are the same. Moreover, the microstructure is very uniform across the width of the substrate. Because the microstructure is related with mechanical properties of samples, this uniform microstructure is desirable for reliability and performance in production situations [7].Figure 4 shows the microstructure of welding sample. The first layer is a regular columnar structure across the width of the substrate, but the microstructure changes rapidly with depth. And later layers become non-uniform across the width and change significantly with positions.284Figure 4.

And later layers become non-uniform across the width and change significantly with positions.284Figure 4. The microstructure sketch of welding sample285This non-uniform microstructure is caused by different cooling rate at different areas. Thenature of the manual arc welding process induces some randomness into the weld. This makes the reproducibility and performance of the repair unsatisfactory.Comparison Between Deposition on Substrates with Different StepsFigure 5 shows the interfacial area of the sample deposited on the substrate with 0.1mm step. The dark area is the substrate, and the grey area is the deposition. The big crack between deposition and substrate clearly indicates that these two sections were not bonded well and this is a failure deposition.depositionFigure 5.

The big crack between deposition and substrate clearly indicates that these two sections were not bonded well and this is a failure deposition.depositionFigure 5. Sample deposited on the substrate with 0.1mm stepIn the same way, figure 6 and figure 7 show the interfacial area of the sample deposited on the substrate with 0.05mm and 0.01 steps respectively. The crack observed from figure 6 elucidates the deposition on this type of substrate is still not success, while figure 7 shows that the step is melted away and there is no defect found from the bonding.It can be concluded that this hybrid repair system is able to successfully deposit on the uneven substrate whose height difference is smaller than 0.01mm. With the substrate whose height difference is bigger than 0.05mm, the tool track and deposit parameters need to be adjusted to make a good deposition.

With the substrate whose height difference is bigger than 0.05mm, the tool track and deposit parameters need to be adjusted to make a good deposition. Some more works are needed to explore deposit ability on the substrates with step height from 0.01mm to 0.05mm.286substrateFigure 6. Sample deposited on the substrate with 0.05mm stepFigure 7. Sample deposited on the substrate with 0.01mm step287Hardness and Tensile StrengthThe Rockwell C-Scale hardness was measured for deposition and welding samples. In addition, because of the different cooling rate, the hardness changes with depth. Therefore, the hardness is listed with different heights which were measured from the interface to the cladding section and with 5mm interval.Table 3.

Therefore, the hardness is listed with different heights which were measured from the interface to the cladding section and with 5mm interval.Table 3. The Hardness of Deposition and Welding Sample (HRC) Deposition 53.7 40.0 56.2 50.0Position 1 Position 2 Position 3 AverageWelding 43.0 48.1 44.5 45.2From the results showed in Table 3, it is clear that the average hardness of laser deposition samples is higher than that of welding samples, however, the hardness of position 2 deposition is the lowest. The hardness of material can be approximately transferred to its tensile strength [8] and the sample will be broken from its weakest section, i.e. deposition sample will be broken from position 2 and welding sample broken from position 1. Thus, the tensile strength for the samples can be estimated as Table 4 listed.Table 4.

deposition sample will be broken from position 2 and welding sample broken from position 1. Thus, the tensile strength for the samples can be estimated as Table 4 listed.Table 4. Tensile Strength Estimated from Hardness (ksi) Deposition 184.0Tensile Strength (Approx.)Welding 201.6For the weakest section, the deposition sample is weaker than welding sample, and that is why its fracture surface can present more ductile feature. In this way, more works such as adjusting the laser deposition process and applying an effective heat treatment are needed to improve the mechanical properties of this section.ConclusionThe hybrid laser metal deposition and machining process can be used to repair casting dies.And the microstructure of laser deposition sample is much more uniform in both the thickness and width directions than that of TIG welding sample. This indicates that more uniform mechanical properties are possible with the laser deposit process.

This indicates that more uniform mechanical properties are possible with the laser deposit process. The average hardness of deposition sample is higher than that of welding sample, but the fact that deposition is weaker at the weakest area will lead to a lower tensile strength. This problem could be solved by transferring the ductility to the hardness via heat treatment and optimizing the deposit process. So far, the hybrid repair system can deposit on the uneven surface whose height difference is smaller than 0.01mm, and the process needs to be adjusted if the height difference is larger than2880.05mm.AcknowledgementThis research was supported by the National Science Foundation Grant NumberDMI-9871185, the grant from the U.S. Air Force Research Laboratory contract # FA8650-04-C-5704, and UMR Intelligent Systems Center. Their support is greatly appreciated. The author also wishes to express his gratitude to: Spartan Light Metal Products Inc. Mr.

Their support is greatly appreciated. The author also wishes to express his gratitude to: Spartan Light Metal Products Inc. Mr. Max Vath, Rolla Technology Institute Mr. Jeff Thomas, Interdisciplinary Engineering at University of Missouri-RollaReferences[1] Eiamsa-ard, Kunnayut , Hari Janardanan Nair, Lan Ren, Jianzhong Ruan, Todd Sparks, and Frank W. Liou, "Part Repair using a Hybrid Manufacturing System", Proceedings of the Sixteenth Annual Solid Freeform Fabrication Symposium, Austin, TX, August 1-3, 2005. [2] Deepak G. Bhat, Francis F. Williams, “Hard Wear Resistant Coatings for Die-Casting Diesby an Advanced Filtered Cathodic Arc Deposition Process”[3] R.E. Smallman, R.J. Bishop, Modern Physical Metallurgy & Materials Engineering,Butterworth-Heinemann, 1999[4] Jianzhong Ruan, Kunnayut Eiamsa-ard, and F.W. Liou, “Automatic Process Planning and Toolpath Generation of a Multiaxis Hybrid Manufacturing System”, Journal of Manufacturing Processes Vol. 7/No.

aluation of Microstructure and Properties for Multi-Materials Laser Densification of Dental RestorationsXiaoxuan Li*, James E. Crocker*, Erik Geiss* Leon L. Shaw*, Harris L.Marcus*, and Thomas B. Cameron**Institute of Materials Science University of Connecticut Storrs, CT**Degussa-Ney Dental Inc. Bloomfield, CTABSTRACTTraditional dental restorations are produced by the porcelain-fused-to-metal (PFM) process, in which a dental restoration is cast from a metallic alloy and then covered with dental porcelains by several firing processes, which is both labor intensive and expensive. In this paper, the feasibility of dental restorations is investigated using a multi-materials laser densification (MMLD) process. To evaluate the effectiveness of the MMLD process, nickel powders and commercial dental porcelain powders are laser densified using YAG and CO2 lasers respectively. Effects of processing parameters, e.g.

Effects of processing parameters, e.g. laser scanning rate and target temperature, are evaluated and the microstructure of processed nickel and porcelain materials are characterized for the optimization of laser densification. Results indicate that densities of laser processed nickel and dental porcelain are strongly dependent of processing parameters. Fully dense layers are achievable with proper processing conditions.KEY WORDS: Multi-materials laser densification, dental restorations, solid freeform fabrication.INTRODUCTIONThere are between 10,000 and 15,000 dental laboratories in the US and a majority of these laboratories use porcelain-fused-to-metal (PFM) restoration for permanent fixed prosthodontics. There are several steps in PFM restoration as shown in Fig. 1, which shows that the traditional PFM process is both time consuming and labor intensive work.

There are several steps in PFM restoration as shown in Fig. 1, which shows that the traditional PFM process is both time consuming and labor intensive work. To address this problem, we have investigated the feasibility of using solid freeform fabrication (SFF) to construct dental restorations. The approach we have proposed is point-by-point and layer-by- layer freeform dental restoration through laser densification of multiple dental alloys and porcelains, which will be termed hereafter as multi-materials laser densification (MMLD). MMLD shares many attributes with selective laser sintering (SLS) [1] and direct metal deposition (DMD) [2]. In SLS, a layer of powder is delivered on top of the substrate (or a previously processed layer) before a laser beam scans the powder layer selectively [1].

In SLS, a layer of powder is delivered on top of the substrate (or a previously processed layer) before a laser beam scans the powder layer selectively [1]. In DMD, a laser generates a melt pool on a substrate (or a previously processed layer) while the feedstock material is introduced into the laser-created pool either as powder or as wirefeed that melts and forms a metallurgical bond with the substrate [2].

The MMLD process to be developed will deal159with multiple materials, and further, dental powders will be delieved and laser densified point- by-point because of the multiple materials involved in almost every single layer.Wax pattern produced from positive 3-D model of patient’s teethSolid investment formed by wax pattern using a ceramic slurryFiring of opaque porcelain powders on dental alloy patternRemove of loose investment, grinding and oxidation of dental alloy patternDental alloy formed by casting into ceramic investmentFiring of translucent dentine porcelain powdersFiring of transparent incisal porcelain powdersFig.

1 Traditional PFM restoration processAs the first step in developing the MMLD process, the current study has focused on the laser densification condition for achieving fully dense bodies from dental alloy and porcelain powders.EXPERIMENTALStarting Materials - Since nickel has been used as the major constituent (79%) in a commercial dental alloy, produced by Degussa-Ney Dental Inc. and widely used in the PFM restoration, a commercial nickel powder was selected as the starting metal for laser densification investigation. The physical properties of the nickel powder used are shown in Table 1. A commercial dental porcelain powder provided by Degussa-Ney Dental Inc. was chosen as the starting porcelain for investigation.

A commercial dental porcelain powder provided by Degussa-Ney Dental Inc. was chosen as the starting porcelain for investigation. Table 2 shows some properties of the selected dental porcelain.Laser Densification Process - The laser densification was carried out in an integrated SFF system composed of a laser, a process chamber, powder delivery system, 3-D laser beam scanning control stage, vacuum and atmosphere control system, observation and recording system, and closed-loop temperature control system. Detailed description of the SFF apparatus can be found in Ref.[3]. The nickel layers were fabricated using a Nd: YAG laser in a continuous wave mode with a maximum output power of 150W and a wavelength of 1.064 microns. The dental porcelain powder layers were manufactured by a continuous wave, maximum power 50W CO2 laser with a wavelength of 10.6 microns.

The dental porcelain powder layers were manufactured by a continuous wave, maximum power 50W CO2 laser with a wavelength of 10.6 microns. The laser beam was steered into the vacuum chamber with the pressure of up to 60 millitorrs by a series of motion-controlled mirrors. The densification was performed through scanning a thick powder bed (about 10mm thick) in a linear pattern. During densification the average surface temperature of the powder under the laser beam was monitored by an emissivity-measuring infrared pyrometer and maintained by a closed-loop control system to match a user-defined target temperature.Table 1. Nickel powder physical propertiesSize -300 meshShape SphericalApparent density 3.6 g/cm3Melting Temperature 14530CPurity 99.9%Coefficient of Thermal Expansion 13.3 x 10-6 K-1160Table 2.

Properties of dental porcelain powderSize 1~50 micronsShape irregularGlass Transition Temperature 6500CCoefficient of Thermal Expansion 14.4~14.9 x 10-6 K-1Two laser processing parameters, i.e. the scanning rate and the target temperature, were controlled to evaluate the effect of processing conditions on the density of processed nickel and dental porcelain bodies. The combination of the various processing parameters used to form the primitive samples is summarized in Tables 3 and 4.Characterization of Microstructure - The primitive samples obtained from these experiments were mounted using an epoxy resin, cut along the direction perpendicular to the scanning direction using a diamond saw, ground and finish polished using an 1-micron diamond suspension. For densified nickel bodies, the samples were examined using an optical microscope directly.

For densified nickel bodies, the samples were examined using an optical microscope directly. For dental porcelain bodies, the samples were etched in 1% HF for 20 seconds, rinsed with water, and then carbon sputter-coated before examination using an environmental scanning electron microscope (PHILIPS ESEM 2020).No.A-1 A-2 A-3 A-4 A-5 B-1 B-2 B-3 B-4 B-5 C-1 C-2 C-3 C-4 C-5Table 3. Nickel powder laser densification (150W YAG) parameters Target Surface Temperature (C) 1200 1200 1200 1200 1200 1300 1300 1300 1300 1300 1400 1400 1400 1400 1400Scanning Rate (micron/second) 2.0 3.8 5.7 7.6 10.0 2.0 3.8 5.7 7.6 10.0 2.0 3.8 5.7 7.6 10.0Beam Diameter (mm) 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2Number of Scanning Pass one one one one one one one one one one one one one one oneTable 4.

Dental porcelain powder laser densification (50W CO2) parametersNo.D-1 D-2 D-3 E-1 E-2 E-3Target Temperature (0C) 800 800 800 900 900 900Scanning Rate (micron/second) 8 16 32 8 16 32Beam Diameter (mm) 1 1 1 1 1 1Number of Scanning Pass one one one one one one161RESULTSMicrostructure of Laser Processed Nickel Bodies - The microstructure of processed nickel in all the trials is similar and can be divided into two zones, as shown in Fig. 2(a). Zone I is almost fully dense with little or no porosity (see Fig. 2b). This zone is formed via melting and solidification. Zone II is a transition region between the dense body and the initial loose powder bed, and has a microstructure with a density gradient. Presumably, this zone is formed partly through incomplete infiltration of the liquid nickel from Zone I and partly through localized powder sintering.

Presumably, this zone is formed partly through incomplete infiltration of the liquid nickel from Zone I and partly through localized powder sintering. The temperature gradient present during laser scanning leads to a gradient in the density of the powder sintered region in Zone II. Viewing from the laser scan direction, Zone I has a V-shape profile, as shown in Fig. 2(a). The investigation to identify the mechanism(s) responsible for the formation of V-shape profile is currently under way.Fig. 2 The typical microstructure of a laser densified nickel layer: (a) the overall shape of the densified nickel with Zone I and II, (b) a high magnification of Zone I, (c) a high magnification of Zone II, and (d) the powder area far away from Zone I. Many powder particles in (d) have been pulled out during polishing, resulting in a very low apparent compact density.

Many powder particles in (d) have been pulled out during polishing, resulting in a very low apparent compact density. The scanning conditions are: target temperature 14000C, scanning rate 10 m m/s, and beam diameter 2 mm.Microstructure of Laser Processed Porcelain Bodies - Dental PFM feldspathic porcelains contain crystalline leucite, a potassium aluminum silicate mineral (K2O.Al2O3.4SiO2)162with a high coefficient of thermal expansion (CTE) in order to elevate the bulk CTE of the dental porcelain to a level compatible with PFM dental alloys [4]. However, the large mismatch between the CTE of leucite and surrounding glass matrix of the porcelain creates thermal mismatch stresses that can cause microcracking around the leucite particles during the PFM firing. Fig. 3 shows a typical microstructure of a fully dense dental porcelain body after traditional firing. From Fig. 3(a) it can be seen that the overall microstructure is composed of two phases: feldspathic glassy matrix (i.e.

From Fig. 3(a) it can be seen that the overall microstructure is composed of two phases: feldspathic glassy matrix (i.e. the featureless region except microcracks) and colonies of leucite crystals. Fig. 3(b) and (c) show the detail of a cluster of leucite particles with different magnifications, and on the top and bottom of Fig. 3(b) there are microcracks. Fig. 3(d) reveals the morphology of the glassy matrix with several microcracks. When feldspathic dental procelain is cooled, the leucite crystals contract more than the surrounding glassy matrix, leading to the development of radial tensile stresses around the leucite particles and thus microcracks within and around the crystals, as shown in Fig. 3(b).The microstructure of a laser densified dental porcelain body is shown in Fig. 4. It can be seen that the top region is denser than the bottom region because of the temperature gradient. The fully dense area has almost the same microstructure as that shown in Fig. 3(b) and (c).

4. It can be seen that the top region is denser than the bottom region because of the temperature gradient. The fully dense area has almost the same microstructure as that shown in Fig. 3(b) and (c). Finely dispersed small crystalline leucite particles lead to reduced radial tensile stresses around the leucite particles; thus less microcracks are found in Fig. 4(b) and (c) than in Fig. 3.163Fig. 3 The typical microstructure of dental porcelain body after traditional firing: (a) the general view; (b) and (c) details of a cluster of crystalline leucite; and (d) the microstructure of feldspathic glassy matrix.164Fig. 4 Microstructure of laser densified dental porcelain bodies: (a) overall profile of a laser processed layer, (b) and (c) colonies of different particles magnifications.

4 Microstructure of laser densified dental porcelain bodies: (a) overall profile of a laser processed layer, (b) and (c) colonies of different particles magnifications. The processing parameters: target temperature 9000C, scanning speed 32 micron/sec.leuciteatEffects of Processing Parameters on Densified Nickel Bodies - The size of Zone I is found to be dependent of the scanning rate and target temperature. For a given target temperature, the width and depth of the densified nickel layer decrease with an increase in the scanning rate. The relationship is shown in Fig. 5. For a given scanning rate the dimension of the densified body increases with increasing the target temperature, as shown in Fig. 6.

The relationship is shown in Fig. 5. For a given scanning rate the dimension of the densified body increases with increasing the target temperature, as shown in Fig. 6. These experimental results are consistent with the expectation on the basis of changes in the heating time and energy input with variation of the scanning rate and target temperature.6655443322110024681012Beam scanning rate, micron/secondFig. 5 The depth and width of laser densified nickel bodies as a function of the scanning rate with a target temperature 14000C.16544.53.5433.52.5232.521.51.5110.50.50 11001200 1400 1300 Target Surface Temperature, C0 1500Fig.

6 The depth and width of laser densified nickel bodies as a function of the target temperature at a scanning rate of 7.6 micron/second.Effects of Processing Parameters on Densified Dental Porcelain Bodies - The relationship between the dimension of laser densified dental porcelain bodies and the processing parameters are similar to that of nickel bodies, as shown in Fig. 7 and Fig. 8.54.54.5443.53.5332.522.521.51.5110.50.50 7007508008509009500 1000Target Surface Temperature, CFig. 7 The dimension of laser densified dental porcelain bodies as a function of target temperature at a scanning rate of 16 micron/second.16632.52.5221.51.5110.50.5005913172125293337Beam scanning rate, micron/secondFig.

8 The depth and width of laser densified dental porcelain bodies as a function of the scanning rate at a target temperature of 8000C.CONCLUSIONS(a) Morphology and size of the laser densified nickel and dental porcelain bodies are dependenton scanning conditions.(b) The size of crystalline leucite particles is smaller in the laser processed porcelain bodies than that processed through traditional firing. Accordingly, less microcracks are found in the laser processed porcelain bodies.(c) Fully dense nickel body is achievable under appropriate processing conditions.Acknowledgement – The authors gratefully acknowledge financial support provided bythe National Science Foundation under Grant No: DMI-9908249.REFERENCES1) J. J. Beaman, J. W. Barlow, D. L. Bourell, R. H. Crawford, H. L. Marcus and K. P. McAlea, “Solid Freeform Fabrication: A New Direction in Manufacturing”, Kluwer Academic Publishers, MA (1997).2) J. Mazumder, J. Choi, K. Nagarathnam, J. Koch and D.

 Evaluation of Non-Stochastic Lattice Structures Fabricated Via Electron Beam Melting.O. Cansizoglu, D. Cormier, O. Harrysson, H. West, and T. Mahale North Carolina State University Edward P. Fitts Department of Industrial and Systems Engineering Raleigh, NC 27695-7906AbstractMetal foam structures have many applications and can be used as structural supports, heat exchangers, shock absorbers, and implant materials. Stochastic metal foams having different cell sizes and densities have been commercially available for a number of years. This paper addresses a different type of foams which are known as non-stochastic foams, or lattice structures. These foams have a well defined repeating unit cell structure rather than the random cell structure in commercially available stochastic foams. The paper reports on preliminary research on the fabrication of non-stochastic Ti-6Al-4V alloy foams using the Electron Beam Melting process.

The paper reports on preliminary research on the fabrication of non-stochastic Ti-6Al-4V alloy foams using the Electron Beam Melting process. Behavior of the structures in compression, bending, and low cycle repeating load tests are discussed, and recommendations about cell geometry and processing conditions are made.Introduction Stochastic metal foams have been commercially available for some time, and are used in a wide variety of applications. These foams can have either open or closed cell structures. Closed cell foams are typically produced by adding gas directly to melt, adding gas releasing blowing agents or causing precipitation of gas bubbles (Banhard, 2000). The process is controlled to achieve the desired pore size, pore shape, and pore distributions, which in turn influence the material properties (see Brothers and Dunand, 2006). Combustion synthesis has also been used to produce metallic foams.

Combustion synthesis has also been used to produce metallic foams. Reactants are mixed, cold pressed and placed in an inert atmosphere and ignited to start the thermal explosion mode or self-propagating high thermal synthesis. Zhang et al. (2001) used combustion synthesis to produce porous surfaces having bone ingrowths without immune response. Powder metallurgy (PM) is a common technique to produce cellular structures where porosity is a natural result of the process. Porosity levels may also be increased by adding blowing agents. The powder mixture is compacted to get a semi finished product. This is followed by heat treatment at near melting temperature of the matrix material. The blowing agent releases gas during the heat treatment. These methods are suitable for Al, Zn, and Mg foams, however Ti is unsuitable due to its high melting temperature and reactivity. Conventional sintering of titanium alloy powders requires sustained high temperature and vacuum. Wen et al.

Conventional sintering of titanium alloy powders requires sustained high temperature and vacuum. Wen et al. (2002) fabricated titanium foams using PM including space holder particles where the resulting densities varied from 50% to 65%. Baumeister et al. (1997) used PM to manufacture aluminum foams. Metallic powders were mixed with foaming agents and compacted to semi finished products. Li et al. (2006) used polyurethane struts produced through RP as a mold and coated them with titanium (Ti-6Al-4V) slurry. After the polyurethane was burned off, the powder was sintered resulting in a porous titanium structure with a porosity of 88% and a compressive strength of 10 MPa. Medlin et al. (2004) report on a method for making tantalum Trabecular Metal®. Carbon skeletons are first obtained by pyrolizing thermosetting polymer foams with repeating dodecahedron structure.

Medlin et al. (2004) report on a method for making tantalum Trabecular Metal®. Carbon skeletons are first obtained by pyrolizing thermosetting polymer foams with repeating dodecahedron structure. Pore size and density of the original polyurethane foam affects209the pore size and mechanical properties of the Trabecular Metal®. A 50μm tantalum film is then deposited onto this carbon skeleton via chemical vapor deposition with a 50μm coating thickness. This process results in a favorable microstructure resulting from proper orientation during deposition and crystallographic growth. Trabecular Metal® has been reported to have superior mechanical properties with an elastic modulus of 3Gpa. Wadley et al. (2003) discussed manufacturing methods of periodic cellular metals.

Trabecular Metal® has been reported to have superior mechanical properties with an elastic modulus of 3Gpa. Wadley et al. (2003) discussed manufacturing methods of periodic cellular metals. Metallic sandwich panels with periodic open cell cores were found to possess better load sustaining capabilities than stochastic foams.Figure 1 Stochastic Copper FoamSFF processes have also been used to fabricate repeating mesh structures. 3D mesh structures in stainless steel and cobalt-chrome were built with Selective Laser Melting (SLM) technology and resulted in nearly 90% reduction in weight (www.mcpgroup.com).The remainder of this paper describes the testing and evaluation of non-stochastic lattice structures produced via the Electron Beam Melting (EBM) process. There were several objectives associated with this preliminary research. The first was to experiment with process settings needed to fabricate these mesh structures.

There were several objectives associated with this preliminary research. The first was to experiment with process settings needed to fabricate these mesh structures. The second was to determine limits on the achievable feature sizes and cell densities. A third objective was to select one cell geometry and gather preliminary material properties.Experimental Procedure Preliminary experimentation started with determining process settings that would produce reasonably cohesive cellular structures. As the objective to this point has been to study the bounds of what can feasibly be made via the EBM process, no attempt has yet been made to optimize the process settings. Starting with initial settings suggested by Arcam, experimentation led to the following melting procedure for lattices. The entire powder bed is first scanned 20 times at 10,000 mm/sec with the beam power gradually ramping up from 6W on the first scan to 1200W on the final scan.

The entire powder bed is first scanned 20 times at 10,000 mm/sec with the beam power gradually ramping up from 6W on the first scan to 1200W on the final scan. The lattice geometry for a given layer is then melted at 180 mm/sec with a beam power of 120W. These process settings were used for all lattices described in this paper.During initial experimentation, it was observed that lattices whose struts are oriented at an angle of less than roughly 35-40(cid:113) with respect to the build plane had little or no structural integrity.210Figure 2 illustrates the problem. The default layer thickness with the EBM process is 0.1 mm. At this layer thickness, it is clear that continuity between the layers becomes a problem as the strut angle relative to the built plate decreases.

At this layer thickness, it is clear that continuity between the layers becomes a problem as the strut angle relative to the built plate decreases. Theoretically speaking, the overlapping distance between adjacent 0.1 mm thick layers can be expressed as L1-L2, where L1=T/sin((cid:302)), L2 = 0.1mm/tan((cid:302)), T is the strut thickness, and (cid:302) is the angle of the strut with the build plane.Figure 2 Overlapping Area Between Successive LayersFigure 3(a) shows a relatively cohesive strut, whereas Figure 3(b) shows a region of relatively poor interlayer bonding resulting from the processing parameters used.

These micrographs help to illustrate that the research presented here is very much a starting point, and that further process optimization aimed at improving strut quality is expected to yield improved material properties.(a)(b)Figure 3 Range of Strut QualitiesAlthough numerous different cell geometries have been built, the material properties presented in the remainder of this paper pertain to the hexagonal unit cell that was modeled in Solidworks. Different views of a 3-D array of these unit cells is shown in Figure 4. Unit cells with different overall cell sizes were designed, with each cell having connecting struts of the same 0.7 mm thickness. By maintaining constant beam thickness while increasing or decreasing the cell size, it was possible to compare lattice material properties as a function of overall density. Magics (Materialise) was used to prepare the build files for fabrication on an Arcam EBM S12 (Arcam, Sweden).

Magics (Materialise) was used to prepare the build files for fabrication on an Arcam EBM S12 (Arcam, Sweden). Slicing of the files was done at 0.1mm layer thickness. The steel start plate was heated up to 750°C before the first layer of powder was melted using the process conditions previously211described. All specimens describe in this paper were built from Ti-6Al-4V powder. Material testing was conducted using an ATS 1605C universal tester. Batches of compression specimens were approximately 25 mm x 25 mm x 25 mm using unit cell dimension of 4 mm, 5 mm, and 6 mm respectively. The specimens were compressed between steel plates at the rate of 5 mm/min. Some samples were also crushed to observe the failure mode. Three-point bending tests were also conducted in same machine using a 50.8 mm span. Length of the bending specimens was approximately 60 mm. The length:thickness ratio was approximately 4 for all bending specimens.

Length of the bending specimens was approximately 60 mm. The length:thickness ratio was approximately 4 for all bending specimens. The structure modulus and compressive strength were calculated according to the head-displacement reading from compression tests. Specimens were tested in the direction parallel to build direction and also in perpendicular to build direction respectively. These are referred to as the XY and Z orientations respectively.(a)(b)(c)Figure 4 - Top, Front, and Dimetric Views of Hexagonal Lattice StructureHead displacements were used to calculate the flexural modulus based on the equationE(cid:32)(3 / LyP ) I 48,where E is modulus (MPa), P is load (N), y is deflection (mm), L is loading span (mm), and I is the overall cross sectional moment of inertia (Kg-m2).Results and Discussion Figure 5 shows a variety of non-stochastic cellular structures that were built in a single run on the EBM machine.

The same hexagonal cell geometry was used in each case, but replications of structures having 4mm, 5mm, and 6mm unit cell dimensions were built.212Figure 5 Compression Specimens with 4 mm, 5 mm, and 6 mm Unit Cell DimensionsFigure 6 shows compression test results for five specimens with 4-mm unit cell dimensions. The XY1, XY2, and XY3 specimens were compressed in a direction parallel to the build direction. The Z1 and Z2 specimens were compressed perpendicular to the build direction. As can be seen from the data, parts loaded parallel to the build direction were stiffer and stronger, but required less work to failure. This trend was also observed for the 5-mm and 6-mm unit cell structures.Figure 6 Compression Test Results for 4mm structuresTable 1 summarizes compression test results for the 4-mm structures. Structures built with a 4- mm unit cell dimension have a relative density of less than 11% of solid titanium alloy.

Structures built with a 4- mm unit cell dimension have a relative density of less than 11% of solid titanium alloy. The213average measured compressive modulus for specimens loaded parallel to the build direction was 211 MPa.Specimen 4mm XY 1Peak Load (N) 4388.08Area (mm2) 597.74Height (mm) 25.5Compressive strength P/A (MPa) 7.341P/A average (MPa)Relative densityEc (MPa) 202.4034mm XY 25472.28585.6423.79.3448.78211.1314mm XY 35658.57585.6423.79.6620.11220.2834mm Z 14mm Z 23754.502995.16561.69561.6924.224.26.6845.3326.0199.27579.072Table 1: 4-mm Structure Compression Test ResultsAs a means of estimating compressive strength of modulus for structures built with different cell densities, the appropriate formulas from Gibson and Ashby (1988) were consulted.

The compressive strength scaling constant was calculated according to the formula(cid:86) (cid:32) cC (cid:86) * 1, sc(p ps)5.1using a solid compressive yield strength of 870 MPa for Ti-6Al-4V. The constant C1 was set to 0.25 to estimate compressive strength for compression specimens. The standard modulus scaling formulaECE (cid:32) 2p ( s ps)2was used with Es=110 Gpa. The constant C2 (geometric constants of proportionality) was found to be 0.1656. The modulus and compressive strength results are shown in Figures 7 and 8 with estimations up to 0.2 (i.e. 20%) relative density.214Figure 7 Compressive Strength As a Function of Relative Structure DensityFigure 8 Elastic Modulus as a Function of Relative Structure Density215Following completion of initial compression tests, a set of bars were fabricated for use in 3-point bending tests. A 50.8 mm span was used for all the tests. Loading was done at 0.5 mm/min for all bending specimens.

A 50.8 mm span was used for all the tests. Loading was done at 0.5 mm/min for all bending specimens. As shown in Figure 9, a subset of the specimens had solid 0.7 mm thick skins built on the top and bottom surfaces. The aim was to increase stiffness and "pin" the ends of the cell struts thus preventing early failure of the meshes. Figure 10 shows the experimental setup.Figure 9 Bending Specimens With and Without SkinsFigure 10 Three Point Bend TestFigure 11 shows bending test results for the 4-mm structures. Those lines with "skin" in the label had the 0.7 mm thick top/bottom skins, whereas the others did not. It is abundantly clear from the plots that adding a thin skin to the specimens significantly increased stiffness and strength, and extended the plastic region.216Figure 11 Bending-Load Displacement Plots for 4-mm StructuresFor the sake of comparison, a finite element simulation model mimicking the conditions of the bending tests was created in Solidworks cosmos.

Figure 12 shows a screen dump from the simulation. The FEA model predicts a compressive modulus of 647.2 MPa. The fact that the predicted modulus is off by approximately a factor of 3 is an important point. Many researchers are developing novel cell geometries and developing simulation models to predict material properties for newly proposed designs. While the models may predict how one design will perform relative to another, the SFF processes do not yet produce struts with perfect cross sections as are those used in idealized models.217Figure 12 - FEA Model for 4-mm Bending TestsConclusions and Future Directions This paper has reported on preliminary experiments involving the fabrication of titanium alloy non-stochastic lattice structures. It was determined that the smallest strut dimensions that can be fabricated without process optimization is approximately 0.7 mm in diameter.

It was determined that the smallest strut dimensions that can be fabricated without process optimization is approximately 0.7 mm in diameter. It is furthermore recommended that the strut angle relative to the horizontal build plane not be less than approximately 35(cid:113) when using 0.7 mm diameter struts. At angles shallower than this, the connectivity between layers is low, and material properties are unacceptable.A variety of cellular structures have been built in a variety of cell densities ranging from 5% to 11% that of solid Ti-6Al-4V. It was determined that adding solid skins to the tops and bottoms of these structures perpendicular to the expected loading direction dramatically improves material properties.In terms of future directions, the most obvious need is a more formal process optimization effort to improve strut quality and hence material properties.

Part of this study could involve post processing operations such as heat treating to effectively "heal" some of the crack initiation sites.Many researchers are exploring with different repeating cell geometries. A comprehensive study in which the relative capabilities of each type of geometry are mapped out would be very useful for designers seeking to find suitable cell structures for a given application.A problem for any of the powder-based (laser and e-beam) SFF approaches is powder removal. Most researchers build structures without skins. However, many real world components require covering with solid skins. Another future topic of study will be to look at the material properties of structures built with solid skins that trap the unmelted powder surrounding the lattice. The218density of the component will still be lower than a solid part due to the fact that the unmelted powder has a lower density.

The218density of the component will still be lower than a solid part due to the fact that the unmelted powder has a lower density. The question to be answered is what the tradeoffs are in strength, stiffness, and weight when a solid skin is built over the part.ReferencesBanhart, J., (2000) Manufacturing Routes for Metallic Foams. JOM, 52, 22-27.Baumeister, J., Banhart, J., Weber, M., Aluminium foams for transport industry Materials & Design, Vol. 18,4, 217-220, 1997.Brothers, A. H. and Dunand, D. C. (2006) Amorphous metal foams. Scripta Materialia, 54, 513– 520.Gibson, L. and Ashby, M. (1988) Cellular Solids Structure & Properties, Pergamon Press, NY.http://www.mcp-group.com/rpt/rpttslm_1.htmlLi, J. P., Wijn, J. R., Blitterswijk, C. A. V., and Groot, K., (2006) Porous Ti6Al4V scaffold directly fabricating by rapid prototyping: Preparation and in vitro experiment. Biomaterials, 27, 1223-1235.Medlin, D. Poggie, R.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEVALUATION OF PROCESSING VARIABLES IN LARGE AREA POLYMER SINTERING OF SINGLE LAYER COMPONENTSBy: Justin Nussbaum1, Garrett Craft2, Julie Harmon2, Nathan Crane11Department of Mechanical Engineering, University of South Florida, Tampa, FL 336102Department of Chemistry, University of South Florida, Tampa, FL 33610ABSTRACTAs the additive manufacturing (AM) industry continues to boom, the material palette continues to grow. However, the materials applicable to selective laser sintering (SLS) remains limited. Typically, the scanning laser beam heats each location for milliseconds at a time with a very high heating flux followed by quick cooling. This can create large temperature gradients and high local temperatures.

This can create large temperature gradients and high local temperatures. Many polymers will degrade or fail to densify under these conditions. Due to the economic constraints for point processing, little work has been done to evaluate the sintering process with lower intensities, longer exposure times and larger areas than are typical with conventional SLS. We will report on a new method for simultaneously heating large areas with spatially-controlled heat flux. A demonstration system and test material is presented and characterized. It is then used to evaluate the relationships between heating rates, exposure time, and resulting sintering levels for traditional Nylon 12 powders in single layer parts.INTRODUCTIONAdditive manufacturing (AM) is a rapidly growing field [1]. Besides making prototypes, the applications are expanding into the electronics, medical, aerospace, military and manufacturing industries [1-4].

Besides making prototypes, the applications are expanding into the electronics, medical, aerospace, military and manufacturing industries [1-4]. The requirements from these industries have a large amount of variation depending on the application. AM is a great candidate for meeting these needs because of it’s capability to make parts with complex geometries, in many materials and with a high degree of customization [5].Additive manufacturing (AM) is an industrial manufacturing method used to create 3D components from solid digital computer models. The components are sliced into 2D cross sections of a prescribed thickness. Each 2D cross section is then reproduced from the feedstock material with exceptional accuracy. These 2D cross sections are stacked on top of each other and fused to the layer beneath it, creating components with complex 3D geometry.

These 2D cross sections are stacked on top of each other and fused to the layer beneath it, creating components with complex 3D geometry. The layering nature of AM allows each part to be unique without requiring custom tooling but at the cost of fabrication time [5].Selective laser sintering (SLS) is an additive manufacturing method that produces high quality parts suitable for end-use. Similar to most other additive manufacturing technologies, SLS forms a part by fusing together multiple layers. To form 3D objects with SLS, a focused laser is typically used to impart energy into a preheated powder bed, fusing together the particles in the exposed area. A single area is typically exposed for milliseconds [6-8] at a time as a laser scans a 2D cross section of the part to be created.

A single area is typically exposed for milliseconds [6-8] at a time as a laser scans a 2D cross section of the part to be created. A new layer of powder is then deposited on top of the first and the process repeats as shown in Figure 1 below.793During the few milliseconds of exposure, the powder particles are brought well above their melting temperature which requires high heating rates. In the case of plastics this can cause sensitive materials to degrade as some parts of the material will be significantly hotter than others. Additionally, the best SLS materials are partially crystalline with a well-defined melting point [9, 10]. This allows the powder to be preheated near the melting point without the powder bed fusing together and reduces the difficulty of spreading new uniform layers. Furthermore, the material should have minimal degradation during transient heating or under prolonged preheat in the powder bed.

Furthermore, the material should have minimal degradation during transient heating or under prolonged preheat in the powder bed. Part distortion can also be reduced if the crystallization temperature is well below the melting point [11]. Nylon is currently one of a few select materials which have properties that make it suitable for laser sintering. A vast majority of the materials used for SLS are nylons or nylon-based because of these constraints [12].LaserLaser Scanner (X-Y)Recoating RollerSintered Parte l c y c e Rr e b m a h CRefill Piston (Z)Build Piston (Z)Figure 1 – During the SLS process a laser and laser scanner sinters the powder in the X-Y axis while the recoating roller and refill piston supplies powder for a new layer.Typically, SLS systems are composed of a high power laser, a galvanometer for steering the laser and a powder spreading system contained within a hermetically sealed inert environment.

This leads to relatively expensive and complex systems, putting this out of range for most users [1]. In order to produce parts economically with high spatial resolution, commercial systems generally have a small laser diameter (~0.5 mm) with high scanning speeds (~1-5 m/s) [6-8]. The proposed system is an alternative to SLS which will provide a lower initial cost and is capable of selectively sintering an entire layer of powder in a single exposure using a high power projection system.Effects of ViscosityHigh temperatures and heating rates can degrade some polymeric materials [13, 14]. Chain scission and oxidation can damage polymer chains and reduce performance by affecting chain molecular weight and stability [14]. Additionally, post condensation reactions in the bulk phase at elevated temperatures can cause increases in the molecular weight in unmodified polyamide (nylon).

Additionally, post condensation reactions in the bulk phase at elevated temperatures can cause increases in the molecular weight in unmodified polyamide (nylon). Thus affecting melt viscosity and sintering efficacy.794Polymer sintering typically occurs through a viscous sintering process. It is a two stage process involving the coalescence of powder particles followed by porosity shrinkage and removal. The early phases are described by a simple viscous sintering model developed by Frenkel as seen below in Equation 1 [15]. Liquid phase sintering is driven by surface tension forces which attempts to minimize surface energy by reducing the surface area [10]. As viscosity is a measure of the resistance to flow, in order to fully fuse and coalesce these particles, the polymer must have sufficiently low viscosity to allow the material to flow over the required sintering time [16].

Low strength components with high porosity will result if viscosities are too high or heating times are too short during sintering. During the transition from solid to liquid, the viscosity drastically decreases, allowing the material to flow. In the simple viscous sintering model below, it becomes evident that the zero shear viscosity has a large impact in sintering coalescence by:1/2(1)1/2𝑥𝑥 𝑟𝑟3𝛾𝛾 2𝜂𝜂0𝑟𝑟= ��𝑡𝑡Where x is the joint length of particle coalescence relative to the original radius (r), γ is the surface tension, R is the zero shear viscosity and t is the sintering time. The sintering rate is determined by the material viscosity. Specifically, the zero shear rate which is most relevant because of the relatively low flow velocity experienced during sintering creates minimal shear stresses.

Specifically, the zero shear rate which is most relevant because of the relatively low flow velocity experienced during sintering creates minimal shear stresses. Zero shear viscosity is proportional to the weight averaged molecular weight and is therefore sensitive to the effects of post condensation reactions that increase the molecular weight [17]. However, the low exposure time and high heating rates in SLS create highly transient, position-dependent temperatures. This makes it is difficult to control the viscosity in-situ or tune the process conditions to the material properties. For laser sintering the easiest way of modifying the zero shear viscosity is by adjusting the molecular weight of the material [15]. Furthermore, the molecular weight is an intrinsic material property which increases over time during the preheat and sintering stages, making in-situ control of this property difficult as well.

Furthermore, the molecular weight is an intrinsic material property which increases over time during the preheat and sintering stages, making in-situ control of this property difficult as well. In the proposed projection sintering system the powder is sintered over the course of a few seconds allowing more time for in-situ monitoring and control.𝜂𝜂0Projection SinteringThis paper describes a system capable of sintering single layered parts with a single exposure and evaluates the parts produced by this method. This concept, of a selective single large area exposure can be seen below in Figure 2. The test system uses a well-established recoating method commonly used in SLS machines where a blade is used to spread new uniformly smooth layers. However, in this approach, the laser has been replaced by a high intensity projector.

However, in this approach, the laser has been replaced by a high intensity projector. The off-the-shelf projector in this study has been modified to concentrate 7.3 W of optical power onto a 3.7 cm2 exposure area while also boosting the optical power. The projection system exposes and sinters an entire layer simultaneously over the course of a few seconds. This method allows for795longer exposure times without compromising the overall build time. This may also enable improved properties for sintering of a wider variety of materials.Figure 2 – left) example part created with a single, large area exposure, right) example thermal image of the University of South Florida logo on a nylon 12 powder bed created with projection sintering systemEXPERIMENTAL SETUP FOR PROJECTION SINTERINGTo create parts with this system, powder was evenly distributed onto an aluminum platform. A blade was used to spread a layer of powder 1.5 mm thick.

A blade was used to spread a layer of powder 1.5 mm thick. A thermocouple was placed in the center of the platform halfway through the thickness of the powder. This thermocouple was used to monitor the temperature of the powder to maintain a preheat temperature of 170 °C. A laboratory convection oven was then placed on its back and the platform and powder were placed on a wire platform inside the oven. A borosilicate glass window was placed in the door of the oven for images to be projected onto the preheated powder inside the oven from the projector mounted outside of the heated chamber. The interior of the oven was subjected to standard atmospheric conditions rather than an inert nitrogen environment which is typical of SLS machines. A schematic of the test system is presented below in Figure 3.Figure 3 – In this experimental setup for polymer powders, an aluminum plate holds a 1.5mm layer of nylon powder at a preheat temperature in an oven.

A thermocouple in the center of the powder is used to select the proper preheat temperature while the high power projector exposes and sinters the powder through a borosilicate glass window.796Since this study seeks to understand the interaction between exposure time and intensity these variables must be carefully controlled and their accuracy validated. A white image on a black background was sent to the projector through Microsoft PowerPoint and the corresponding image was projected onto the powder. White pixels represent areas which are receiving the maximum power while black pixels contain the least amount of intensity. Depending on the level of gray scale, varying levels of power can be projected.In order to determine the proper preheat temperature, a differential scanning calorimetry (DSC) study was conducted on the PA2202 with a TA SDT Q600. This DSC study was conducted in an air environment with a heating and cooling rate of 10°C/min.

This DSC study was conducted in an air environment with a heating and cooling rate of 10°C/min. The powder was first dried in roughing vacuum at 70°C for 24 hours. The DSC data (shown below in Figure 4) shows that a phase transition begins around 175°C. This corresponds well with the manufacturer’s reported melting temperature of 172-180°C [18]. On cooling, crystallization begins below 155 °C. To maintain a temperature directly under the melting temperature, a 170°C power preheat temperature was chosen. By preheating the powder to just under the melting temperature, less optical energy is required to melt the powder and reduces the amount of internal stresses which can lead to warping.1.0Working Temperature Range0.5) g / W0Cooling(w o Fl0.5Heatingt a e H1.01.52.020 40600 80 100 120 140 160 180 200Temperature (°C)Figure 4 Differential scanning calorimetry study shows a solid to liquid phase transition at ~175°C.

A preheat temperature was chosen 5°C below this transition at 170°C..Optical PowerAn Optoma X316 off-the-shelf projector was modified to decrease the exposure area from a maximum of 278,709 cm2 to 3.7 cm2. Since the maximum optical power (W) is fixed and independent of screen size, decreasing the screen size increases the power density (W/cm2). Additionally, the optical power was boosted from 1.8 W to 7.3 W, measured from a fully white projected image on a Thorlabs S310C thermal power sensor. Overall, this provides a power density increase from 6.5x10-6 W/cm2 to 1.97 W/cm2, an increase of more than 30,000%.797The projector’s optical power is provided by a 190 W ultra-high pressure mercury vapor (UHP) bulb. UHP bulbs produce a broad spectrum of light that ranges from the ultra violet (UV) to infrared (IR) spectrum. A majority of the energy from UHP bulbs are contained within the visible spectrum (400-800 nm) and produces a mostly white light.

A majority of the energy from UHP bulbs are contained within the visible spectrum (400-800 nm) and produces a mostly white light. The absorption of visible light is strongly connected to the color of the absorbing material. As such, white powders reflect most of the visible light away from the surface, absorbing only a small portion of the optical energy from the light. A dark colored powder absorbs a much larger quantity of energy from the light. In order to maximize the absorbed energy from the light, a dark gray nylon 12 powder (PA2202) produced by EOS was selected for testing. This differs from SLS systems which use a CO2 laser operating in the IR spectrum at 10.6 μm, where energy absorbance is not heavily color dependent as it is with visible light. Upon melting, the PA2202 powder becomes evenly black as the black pigment (carbon black) evenly disperses through the component.

Upon melting, the PA2202 powder becomes evenly black as the black pigment (carbon black) evenly disperses through the component. This material is manufactured for the SLS industry and is commonly used in SLS systems to create black components.The degree of sintering is highly influenced by the amount of time the powder is exposed to high intensity light. A high degree of timing accuracy is required by the system in order to create reproducible results. The timing accuracy of the system was measured by projecting an image onto a high speed photodiode (OPT101 produced by Texas Instruments) which also internally contains a transimpedence amplifier. The photodiode was connected to an Agilent Technologies Mixed Signal Oscilloscope (MSO6012A) capable of 100 MHz sampling frequencies. An image was displayed then removed after a specified animation time and the photodiode output was recorded. The system occasionally produced images for less or more than the specified time.

An image was displayed then removed after a specified animation time and the photodiode output was recorded. The system occasionally produced images for less or more than the specified time. This amount of error lead to a timing inaccuracy of 1ms ± 27ms. Since exposure times ranged from 1-15s during testing, this amount of error in the exposure time was considered negligible.One issue identified from using PowerPoint shapes was that varying levels of grayscale were not linearly proportional to the projected power. By using a Thorlabs S310C thermal power meter, varying levels of grayscale were projected and the power contained within the light was measured by allowing the power meter to reach steady state. This test was conducted three times and the results are shown below in Figure 5.

This test was conducted three times and the results are shown below in Figure 5. By fitting a regression line to the data, the projected power could be calibrated.798Optical Power100%90%80%y = -15.6344x6 + 58.2042x5 - 85.2702x4 + 60.7607x3 - 20.9823x2 + 3.8796x + 0.0404 R² = 0.997770%)%(s s e n t h g i r B60%50%40%Measured30%20%10%0%0%20%40%60%80%100%Measured Power (%)Figure 5 – Optical power as measured by thermal power sensor relative to the desired output capability was mapped to different levels of brightness. A brightness of 100% indicates the highest power and is completely white, while an intensity of 0% is the lowest intensity and is completely black.As the light is emitted from the projector onto the powder, some of the light is absorbed, some of it is reflected and the rest is transmitted through the power.

Energy from the light is only reflected away from the powder at the powder’s surface and is independent of the thickness of the powder. As the light penetrates into the powder bed, the optical power is attenuated as a portion of the light is absorbed by the powder, transmitting the rest. By measuring the attenuation of the transmitted light through powder layers of varying thickness, the amount of absorbed light can be gauged.To conduct this experiment the TI OPT101 photodiode was also used in this experiment to measure the intensity of the light. It was sputter coated with gold-palladium to darken the front face in order to increase the maximum level of intensity the photodiode can read before becoming saturated. The photodiode was permanently fixed to a glass plate which was then leveled and focused to the projectors image. The photodiode and glass plate were then fixed to a three axis CNC machine where a razor blade was fixed to one of the axis.

The photodiode and glass plate were then fixed to a three axis CNC machine where a razor blade was fixed to one of the axis. A layer of powder was manually deposited on top of the photodiode and glass plate and the CNC machine carried out automated code which moved the powder, glass plate and photodiode under the razor blade, spreading uniform layers of powder at varying thicknesses. After each layer of powder was deposited, the CNC machine positioned the photodiode and glass plate beneath the projector where a completely white uniform image was projected. The projector was lifted to a distance of approximately 0.5m away from the photodiode. This decreased the power density by enlarging the projection image to prevent sintering from occurring and to prevent the photodiode’s output voltage from being saturated.

This decreased the power density by enlarging the projection image to prevent sintering from occurring and to prevent the photodiode’s output voltage from being saturated. The photodiode’s output was then recorded and correlated with varying powder layer thicknesses as seen below in Figure 6.799Light Penetration Depth)100%%(r e w o P g n n a m e Rii80%60%40%20%0%0100200300Layer Thickness (μm)Figure 6 - left) schematic of system for measuring light intensity at varying powder depths, right) optical intensity profile.In order to determine what the minimum layer thickness is, layers of powder were spread with the CNC machine in 25 μm increments. The PA2202 powder which was used for these tests have a mean particle size of 55 μm. The minimum layer thickness which could be repeatably spread without smearing was found to be 100 μm. This means the thinnest layer which can be repeatably spread is about two particles thick.

The minimum layer thickness which could be repeatably spread without smearing was found to be 100 μm. This means the thinnest layer which can be repeatably spread is about two particles thick. Layers which are 100 μm thick are standard in the additive manufacturing industry. It is small enough to provide decent Z resolution but not so small that build times are drastically increased or spreading uniform layers becomes difficult. In this first 100 μm of powder, 65% of the light has been absorbed by the powder and at 200 μm, or 2 typical layers into the powder, only 5% of the light remains. These are promising results as a majority of the optical power is absorbed quickly in the first layer leaving only a small portion of the light to penetrate further into the powder.

These are promising results as a majority of the optical power is absorbed quickly in the first layer leaving only a small portion of the light to penetrate further into the powder. This could be advantageous as the transmitted light which is absorbed by the second layer thickness could aid in fusing new layers of powder together.SINGLE LAYER PART TESTINGTo maximize the strength of any type of part or material, the porosity needs to be kept at a minimum. This allows the bulk of the material to bear and evenly distribute any loading forces. To maximize the load carrying capability, the strength should also be maximized. Since porosity and strength are important characteristics of any material, they should be investigated to determine the optimum manufacturing parameters which will maximize the performance of the part.

In this study, components under different levels of sintering are cross sectioned and viewed under an SEM and are also tested for maximum load carrying capability in a tensile testing machine.SEM Evaluation of Porosity and ThicknessTo create parts with adequate strength, the powder must fully melt and densify during sintering. To analyze the degree of densification, four test coupons were sintered under varying degrees of exposure time at the maximum optical intensity. These test coupons were sintered on a powder bed 1.5mm thick (27 times the mean particle diameter). This case should approximate the sintering of the first layer of a part. They were then submerged in a bath of liquid nitrogen and broken in half to create a brittle fracture with a nearly flat fracture surface.

They were then submerged in a bath of liquid nitrogen and broken in half to create a brittle fracture with a nearly flat fracture surface. These components were800then sputter coated with gold-palladium to create a conductive surface and imaged with a Hitachi S800 SEM as seen below in Figure 7.Exposure times below 1.5s produced parts that were too weak to be handled and are not imaged. As seen in Figure 7a, little particle necking took place during sintering with individual particles clearly visible through the entire part thickness, resulting in a low degree of densification and low part strength. The bottom layers show little evidence of sintering. As the exposure time is increased, the particles reach higher temperatures (and thus lower viscosity) and have longer time to densify as seen in Figure 7b. As the exposure time is further increased, the parts further densify and the thickness of the dense region increases as the molten nylon coalesces with neighboring particles.

As the exposure time is further increased, the parts further densify and the thickness of the dense region increases as the molten nylon coalesces with neighboring particles. At longer exposures, a smooth surface layer is formed with few visible pores. Increasing exposure times increases the thickness of this layer. After 3s the layer is approximately 100 µm thick, which is encouraging for the formation of multilayer parts. However, longer exposures also produce curling of the edges of the parts and is discussed in more detail below.Figure 7 - All components were sintered with 100% intensity and are shown at 100x magnification. The exposed surface is on top in all images. a) 1.5s exposure b) 1.75s exposure c) 2.0s exposure d) 3s exposure.801Tensile TestingIn order to analyze the strength of parts produced with projection sintering technology, single layer tensile test specimens were created using a single exposure.

Various sintering tests were conducted and tested in accordance to ASTM standard D638-10 and the combinations of intensity and time are summarized in Table 1. The tensile specimen geometry is similar to the standard, however it has been scaled down to fit the maximum exposure size and the gauge width was increased to increase the total breaking force. The projected image had an overall length of 22.5mm and a gauge width of 7.3mm. All parts were produced with a single exposure. Thickness was determined by the heating time/intensity combination used. After removal from the powder bed, the back of each specimen was lightly brushed to remove loose non-load bearing powder. The specimen thickness was measured with a dial indicator whose contact tip was replaced with a blunted 27 gauge needle. The blunt needle was used to enable measurement of curled samples. The specimens were measured on the long axis centerline at three locations and averaged.

Each of these tests were conducted on a MTS 810 hydraulic tensile testing machine.Table 1 – Fabrication parameters and results of tensile testing coupons, three specimens were tested at each parameter set.Test ConstantIntensityExposure TimeEnergy DoseIntensity100% 100% 100% 100% 100% 100% 100% 100% 100% 100% 100% 90% 80% 70% 60% 50% 40% 100% 90% 80% 70% 60% 50%Exposure Time (s) 1 1.25 1.5 2 2.5 3 4 5 10 15 3 3 3 3 3 3 3 2 2.22 2.5 2.86 3.33 4Power Density (W/cm2) 1.97 1.97 1.97 1.97 1.97 1.97 1.97 1.97 1.97 1.97 1.97 1.78 1.58 1.38 1.18 0.99 0.79 1.97 1.78 1.58 1.38 1.18 0.99Energy Dose (J) 7.3 9.125 10.95 14.6 18.25 21.9 29.2 36.5 73 109.5 21.90 19.71 17.52 15.33 13.14 10.95 8.76 15 15 15 15 15 15Avg Thickness (μm) 161 ± 15 212 ± 39 220 ± 15 288 ± 15 381 ± 0 483 ± 0 656 ± 26 627 ± 44 1080 ± 13 1325 ± 15 445 ± 38 406 ± 13 322 ± 19 246 ± 15 246 ± 15 191 ± 13 N/A 229 ± 13 237 ± 19 254 ± 13 301 ± 15 296 ± 7 292 ± 0Avg Breaking Force (N) N/A N/A 5.9 ± 3.5 22.1 ± 3.6 35.5 ± 5.4 49.5 ± 3.0 62.2 ± 5.5 53.5 ± 19.3 195.2 ± 9.5 226 ± 27.9 22 ± 10 29.6 ± 5.9 22.7 ± 4.1 13.8 ± 1.8 4.5 ± 1.7 N/A N/A 9.7 ± 5 12.1 ± 1.6 9.6 ± 7.1 2.4 ± 2.1 3.1 ± 0.8 9.8 ± 3802During these tests the components which were sintered for a short period of time remained flat but were also of insufficient strength to be handled.

Few could be extracted from the powder bed without damage. Specimens which were well sintered showed a large degree of curling directly after exposure in the powder bed. This made it difficult to measure the cross sectional width in the gauge length. The forces recorded during tensile testing were not able to be converted to stress values because of this. Since forces do not demonstrate a material property, they are used here only for relative comparison.Three different tensile tests were conducted. For the first set of tests, 30 tensile test bars were created, 3 for each of the 10 test parameters. For these tests, the light intensity was held constant at the maximum value. Exposure times was increased from 1-15s which also increased the overall energy input into each specimen (energy dose). By increasing the exposure time (and energy dose), the powder is allowed longer periods of time to densify.

By increasing the exposure time (and energy dose), the powder is allowed longer periods of time to densify. As the exposure is maintained for longer periods of time, the powder is able to reach higher temperatures, and thus lower viscosities. This allows the material to flow more readily and to coalesce with neighboring particles. Increasing exposure time has the effect of creating thicker, denser and therefore, stronger parts. However, considerable curling was occurring with exposures over 1.5s as seen in Figure 8 below. Since the specimens that were produced with 1.5s exposure times only show a minimal amount of strength, the warping mechanism will need to be determined and corrected in order to create functional 2D or 3D parts. With 10s of exposure or more, the part’s edges warped upwards during sintering, uncovering the powder beneath it which was then also sintered to the part, creating a perimeter.

With 10s of exposure or more, the part’s edges warped upwards during sintering, uncovering the powder beneath it which was then also sintered to the part, creating a perimeter. With a 10s exposure, the time and temperature were sufficient to cause stress relaxation which allowed the part to lay flat again.a)b)c)Figure 8 – Tensile specimens made with the maximum intensity a) 1.5s exposure and 10.95J of energy, b) 2.5s exposure and 18.25J of energy, c) 15s exposure and 109.5J of energy.For the second tests, the exposure time was held constant at 3s and the intensity was decreased from 100% to 40%. Three tensile bars were made for each of the 7 test conditions, for a total 21 tensile bars. The exposure time for this test was chosen because it produced a part of decent strength but produced excessive warping. By decreasing the intensity by 10% for each sample set, the optimal time-intensity setting could be identified.

By decreasing the intensity by 10% for each sample set, the optimal time-intensity setting could be identified. As the intensity was decreased (which also decreased the energy dose) lower strength, thinner parts were produced. The optimal intensity value was found to be at 70% intensity and can be viewed in Figure 10 below. This created a part of decent strength with minimal warping. The lowest possible intensity for this exposure time to produce parts that can be handled was identified to be at 60%. Parts below this value left a darkened discoloration in the powder but crumbled upon handling.803Figure 10 – Tensile specimens made with a 3s exposure, a) 100% intensity with 3J of energy, b) 70% intensity with 2.1J of energy, c) 50% intensity with 1.5J of energy.During the third set of testing, the intensity was decreased from 100% to 50% while the total input energy was maintained.

This required the exposure time to increase as the intensity decreased to maintain the overall energy dose. This had an effect of greatly decreasing the amount of warping but at the cost of reducing the load carrying capability. However, the change in the maximum achievable force and thickness varied much less between tests than with the previous two tests. This indicates that the time and intensity are inversely proportional to some degree. If the intensity or the time must be changed, then the other variable should change in the opposite direction as well to maintain the energy dose. It is evident from the pictures seen below, in Figure 10, that less sintering is occurring in the parts which have been exposed for a longer period of time at a lower intensity. However, a lot of variation was seen in these measurements, especially for the 60% and 70% intensity specimens.

However, a lot of variation was seen in these measurements, especially for the 60% and 70% intensity specimens. This could be cause by a non-ideal situation with the samples during handling, loading or during the sintering process. In order to reduce the variation, a higher energy dose would increase the ease of measurement and make the specimens less likely to be damaged before testing. Eighteen tensile bars were made for this set of testing but still requires more work to fully understand. Conducting this test at higher energy doses could help in understanding this effect.

Eighteen tensile bars were made for this set of testing but still requires more work to fully understand. Conducting this test at higher energy doses could help in understanding this effect. Additionally, adapting Andrew’s number (measure of relative energy density) or the energy melt ratio (comparison of actual to theoretical sintering conditions) to projection sintering could aid in finding the correct parameter set to get high quality parts [19, 20].Figure 9 - Tensile specimens made with the same energy dose, a) 4s exposure and 50% intensity, b) 3.33s exposure and 60% intensity, c) 2s exposure and 100% intensity.804CONCLUSIONA new area based sintering technology and system are created to aid in analyzing the role of time and optical intensity in polymer sintering. This system is capable of sintering entire 2D cross sections with a single quick exposure.

This system is capable of sintering entire 2D cross sections with a single quick exposure. The modified projector used in this study was examined for accuracy in exposure time and calibrated to accurately control the intensity. The nylon 12 powder was also characterized to determine melt temperature and visible light absorption. This understanding is critical in the development of the new proposed method of sintering.Projection based sintering can potentially expand the capable materials in the sintering field beyond the commonly used nylon. This system could aid in the understanding of sintering kinetics outside the realm of high heating rates. It is also the first step in creating a system capable of creating 3D printed parts with this technology. This large area sintering technology is modular and can be scaled to industrial sized machines. Additional or higher power projectors could be used to sinter with higher heating rates or over larger areas.

Additional or higher power projectors could be used to sinter with higher heating rates or over larger areas. Fast refresh rates of projectors also allow for dynamic control over spatial temperature distributions within the exposure area. However, many obstacles exist, including warping and correcting for a nonuniform distribution of light across the exposure field. More work still needs to be completed to fully understand the role time and intensity play in sintering kinetics and how it effects the overall part performance. Additionally, future work will have to be conducted to evaluate the effect long exposures have on the minimum feature size that can be obtained with these systems. This preliminary work suggests that energy input per unit area may still be a useful process parameters but may need some adjustments when used for longer processing times.805References[1][2][3][4][5]T. T.

Vlachopoulos, "An experimental study andmodel assessment of polymer sintering," Polymer Engineering & Science, vol. 36, pp. 2198-2207, 1996.[17] C. Tonhauser, D. Wilms, Y. Korth, H. Frey, and C. Friedrich, "Entanglement Transition in Hyperbranched Polyether-Polyols," Macromolecular Rapid Communications, vol. 31, p. 2127, 12/15/ 2010. "PA2202 Material Data Sheet," E. GmbH, Ed., ed, 2009.[18] [19] M. Vasquez, B. Haworth, and N. Hopkinson, "Methods for quantifying the stablesintering region in laser sintered polyamide-12," Polymer Engineering and Science, p. 1230, 2013.[20] D. M. J. Williams, C. Deckard, "Selective Laser Sintering Part Strength as a Function of Andrew Number, Scan Rate and Spot Size," in Solid Free Form Fabrication Symposium, Austin, Texas, 1996.807.

aluation of PROMETAL technique for Application to Dies for Short Run ForgingsBy Kuldeep Agarwal, Deepanker Mathur, Rajiv Shivpuri Department of Industrial, Welding and Systems Engineering The Ohio State University, 1971 Neil Avenue, Columbus, Ohio 43210, USA And John Lembo Extrudehone Corporation 1 Industry Blvd., P.O. Box 1000, Irwin, PA. 15642AbstractManufacturing of hot forging dies required several steps such as acquisition of material block, shaping it for machining, rough machining the cavity, heat treating, finish machining, grinding and polishing. This process takes several months. Consequently for limited number of parts often required in aerospace industry, forging is being replaced by direct machining of parts. If the die lead times (administrative and manufacturing) could be reduced to weeks instead of months, forging process will become viable for short run forgings. This paper evaluates the PROMETAL technique for dies in forging of aluminum alloys.

This paper evaluates the PROMETAL technique for dies in forging of aluminum alloys. This evaluation includes frictional, heat transfer and strength characterization. Isothermal and non-isothermal ring tests together with FEM models are used to determine the interface behavior and its effect on metal flow.IntroductionA typical forging development procedure involves part design, process and material selection, forging design, die and tooling design, tooling manufacture, set up and production, heat treatment, machining of the forging, testing and assembly. The focus of this study is on forging of aluminum alloys.Aluminum alloys are forged into a variety of shapes and types of forgings with a broad range of final part forging design criteria based on the intended application [1].

For a given aluminum alloy forging, the pressure requirements in forging vary widely, depending on the chemical composition of the alloy being forged, the forging process being employed, the forging strain rate, the type of forging being manufactured, the lubrication conditions, and the forging and die temperature.The forging temperatures for most aluminum alloys are in the range of 600-850º F. The flow stress shows a large variation with change in temperature. Thus the forgeability of all aluminum alloys improves with increasing metal temperature. Therefore, unlike some forging processes for carbon and alloy steels, the dies used in virtually all hot-forging processes for aluminum alloys are heated in order to reduce die chill.The first step in the process of die manufacturing for aluminum forging is to select the die block, which satisfies the press space and cavity sinking requirements. The die block is cast from hot working die steels.

The die block is cast from hot working die steels. The block is machined to the CAD models and cavities machined in the soft condition. Then the dies are heat treated and polished after heat treatment. Alternatively, the forging cavities can be sunk by EDM process in the hard condition in the heat treated blocks. For reasonable die lives the cavities sunk by EDM need considerable post-sinking polishing. In376addition, this approach requires the electrodes of the EDM to be machined previously. More recently, the advent of high speed machining has made it possible to directly cut the hardened die blocks. This has reduced die-manufacturing times but increased costs due to enhanced tool wear and high cost of machining centers capable of high spindle speeds.Forgings are dismissed as useful for short run because of the long lead times involved in the manufacture of tooling.

Thus people are forced to revert to machining for short run forgings even when it involved cost and performance penalties. The objective of this study is to develop a rapid tooling design and development system for short run forgings which deploys rapid tool manufacturing techniques together with integrated process and product design approaches. This paper discusses the rapid manufacture of dies.Rapid prototyping of diesPopularity of RP techniques has resulted in large number of techniques being developed. However, only a few have found application for tools and dies. For example, Jacobs and Hilton [2] have written about the inserts made by the technique of Direct AIM (ACES injection molding). Xerox Corporation on an SLA-250 built the inserts with Cibatool SL5170 epoxy resin using ACES build style. They were able to injection mold 100 polystyrene parts just 5 days after the CAD design was complete.

They were able to injection mold 100 polystyrene parts just 5 days after the CAD design was complete. Karapatis and Glardon(1998) [3] presented the various methods of direct rapid tooling with SLS technique. According to them, direct tooling utilizing SLS is possible with metallic powder; polymer, metal mixture (polymer coated metallic powder) or nickel-bronze-copper phosphide powder. King and Tansey (2000) [4] talked about the rapid steel method used to create steel/copper mold inserts for injection molding. . The powder used is stainless steel powder infiltrated with bronze. Griffith and Ensz [5] evaluated properties in LENS process. Simple sample geometries were chosen where the tensile direction is either parallel (H) or perpendicular (V) to the layers. The stainless steel alloys show the greatest effect of the layered deposition, where the strengths are lower for the vertical samples.

The stainless steel alloys show the greatest effect of the layered deposition, where the strengths are lower for the vertical samples. This is most likely due to the stress state condition where the layers are perpendicular to the pull direction and any imperfections will initiate fracture. Agarwala and Osborne [6] talk about rapid prototyped pre forms with HIPping . The ceramic performs are generated by RP processes and temperature and pressure is used to fill the cavity with tool steel. RSP tooling technique [7] gives the material properties close to H-13, but has only been evaluated on injection molds and die casting tools. In this process spray forming of thick deposit of tool steel is done on a pattern to capture the desired shape and texture.Prometal techniqueOne of the promising rapid prototyping techniques which can be used for forging dies is It has been commercialized bythe 3D printing Extrudehone/Prometal.

The main technique of making the prototypes is as follows –technique developed by MIT.(a) The CAD file is sliced into layers and a STL file is generated. (b) Each layer begins with a thin distribution of powder spread over the surface of a powderbed.(c) Using a technology similar to ink-jet printing, a binder material selectively joins particleswhere the object is to be formed.(d) A piston that supports the powder bed and the part-in-progress lowers so that the nextpowder layer can be spread and selectively joined.377This layer-by-layer process repeats until the part is completed. Following a heat treatment, unbound powder is removed, the metal powder is sintered together. This sintered metal powder is then infiltrated with bronze to impart the strength and fill up the pores. The process is shown in fig. 1.Material propertiesThe material properties obtained by the Prometal process are shown and compared with conventional tool steel in table – 1.

The process is shown in fig. 1.Material propertiesThe material properties obtained by the Prometal process are shown and compared with conventional tool steel in table – 1. There are 2 phases in the material – 420 Stainless steel and bronze. This can be seen from the fig. 5.To find out the variation of the hardness in the material in different orientations, a microhardness test was conducted on the sample. A section was cut from the sample and the hardness values (Rockwell C) were taken both along and perpendicular to the orientation. The results are shown in Fig. 6.Initial experimentsA short run forging tool was fabricated within nine business days from receipt of the file. This forging die for Aluminum Precision Products was manufactured with 420 stainless steel metal powder and infiltrated with bronze. The resultant material composition is 60% steel and 40% bronze.

This forging die for Aluminum Precision Products was manufactured with 420 stainless steel metal powder and infiltrated with bronze. The resultant material composition is 60% steel and 40% bronze. The die was used in an isothermal forging application with the forged material being 2XXX series aluminum (fig. 2). The part produced is shown in fig. 3. This proof of concept trial clearly demonstrated the ability to produce a near net shape tool in a rapid time frame.Material characterizationThe first step in the development of dies by 3D printing method is the characterization of material produced by the technique. The issues that come in the study are the friction and heat transfer between the die and workpeice.Friction in metal workingIn metal deformation processes, frictional forces are generated at the interface between the tools and deforming materials by virtue of the workpiece surface extension [9].

These frictional forces have the following modifying effects – (a) The total deformation loads are increased. (b) The internal structure and surface characteristics (surface finish and surface defects) of theproduct are influenced.(c) Wear is produced on the tooling material, thus reducing its useful life. (d) Dimensional variations are produced in the processed material.Because of these effects, friction is considered to be a major variable in metal working operations and must be adequately controlled to optimize processing procedures for economically producing material with the desired geometry and internal structure.

For effective friction control, quantitative data on the effects of lubrication and other processing variables, such as temperature, speed and pressure, are essential.Of the many laboratory tests utilized for friction studies the ring test technique originated by Kunogi [10] and further developed by Male and Cockcroft [11] has the greatest capability for quantitatively measuring friction under normal processing conditions.378The ring test technique involves a simple forging operation performed on a flat ring- shaped specimen, the change in diameter produced by a given amount of compression in the thickness direction is related to the interfacial friction conditions. If friction were equal to zero, the ring would deform in the same way as a solid disk, with each element flowing radially outwards at a rate proportional to its distance from the center.

If friction were equal to zero, the ring would deform in the same way as a solid disk, with each element flowing radially outwards at a rate proportional to its distance from the center. With a small but finite interfacial frictional force, outward flow takes place at a lower rate and, for the same degree of compression, the outside diameter is smaller than with zero friction.Measurement of the final internal diameter of compressed rings provides a particular sensitive means for studying interface friction. Correlation of changes in internal diameter with numerical values of friction can be obtained either by independent calibration or by simulations.Ring compression testThe ring compression test was conducted at OSU. The rings taken were of the standardgeometry ( 6:3:2) –Outer dia – 1”, Inner dia – ½” , Height – 1/3”The platens used to compress the rings were made from Prometal technique and had the orientations of 0˚, 30˚, 60˚, 90˚.

The test was also conducted with die made of H-13. The compressed rings obtained by the test are shown in fig. 7.First, the test was conducted with keeping the dies at room temperature and heating the rings to 800˚F. The friction factor obtained was verified by DEFORM simulations and is shown in table – 2. The test was conducted with and without the use of lubricant (graphite in colloidal solution of water and oil).Because of the steep change in the flow stress of aluminum with change of temperature, the test was conducted again by heating the dies to about 450˚F. The results of this test are shown in table – 2.The final series of test was conducted by polishing the die to a surface roughness of 23microinches. The results of the friction factor obtained are shown in table – 3.After the ring compression test was completed, the cross section of the ring was takenand its microstructure was evaluated.

The results of the friction factor obtained are shown in table – 3.After the ring compression test was completed, the cross section of the ring was takenand its microstructure was evaluated. The microstructures are shown in fig 8.DiscussionsThe ring compression test was performed with rings of standard geometry (6:3:2). The die samples used were of the Prometal material (with 4 different orientations – fig. 4) and H-13. The tests were conducted with and without the use of lubricant and with and without heating the dies. The different orientations of the dies did not show much variation in the friction factor. The friction factor of the material was found to be 0.75 when the dies are at room temperature. This factor reduced to 0.55 when the dies were heated to 450˚F.

The friction factor of the material was found to be 0.75 when the dies are at room temperature. This factor reduced to 0.55 when the dies were heated to 450˚F. Also, this factor reduced to 0.27 (0 degree) and 0.30 (90 degree) when the dies were polished to a surface roughness of 23 or 33µin respectively.The microhardness test shows that the hardness values are different for all orientations and are almost the same along different directions in the same orientation. 0 degree orientation showed the lowest average hardness while 90 degree orientation showed the maximum hardness.The microstructures reveal that with H-13 dies there is a zone of recystallization at the corner while with Prometal dies there is rapid quenching and the aluminum microstructure is homogenous throughout.

Thus it can be deduced that the Prometal dies have a much higher heat transfer coefficient than H-13.379A universal scale down die was designed with the rib-web dimensions and the draft angles that represent a typical aluminum forging (fig. 9). It has been planned to do forging tests on this scale down die and the simulation showing the filling is shown in Fig. 10. The forgings will be done on 0 degree and 90 degree orientation.ConclusionsThe ring compression test showed that the friction at the interface between the die and aluminum under forging conditions is quite high. As mentioned earlier, this can result in problems while actual forging. So, it is necessary to perform some kind of surface treatment on the die made by such a process so that the friction can be reduced which might be suitable for forging applications.The results of the microhardness tests show that the difference in values along a layer are sometimes quite high.

This might be due to the different phases in the material (bronze and SS 420). Thus, the material has different properties in different directions and is thus anisotropic.AcknowledgementsThe authors would like to thank Defense Logistics Agency (DLA) and Advanced Technology Institute (ATI) for providing the opportunity for this study under ProFAST program contract No. SP0103-01-C-0002, M/S-Sikorsky Aircraft for providing the parts required, M/S - Weber Metals for giving inputs and material for forgings. We would also like to thank Mr.

SP0103-01-C-0002, M/S-Sikorsky Aircraft for providing the parts required, M/S - Weber Metals for giving inputs and material for forgings. We would also like to thank Mr. Lloyd Barnhart for his extended support during the experimentation.PropertyHardness Ultimate Strength, Ksi Yield Strength, Ksi Young’s Modulus, Mpsi Elongation Thermal Conductivity Btu.in/ft2.h.F Density, lb/in3Conventional tool steel 51-55 HRC 150 115 30 9 % 169 0.282316 + Bronze60 HRB 59 34 21.5 8 % 51 0.29420 + Bronze26-30 HRC 99 66 21.4 2.30 % 57 0.289Table – 1 Material properties of Prometal technique and conventional tool steelS.No.

2 Die made from Prometal technique Fig. 3 Part used for initial trial of the die30 Degree0 Degree90 DegreeFig. 4 Ring compression dies build direction3810 Degree90 DegreeFig. 5 Microstructure of the samples (50 X)Hardness along different orientations60 50 40 30 20 10 090 degree0 degree30 degree01020Distance from edge (in mm)Fig. 6 Microhardness test resultsReduction – 10-12% 23-25% 40-43% 53-56%Fig. 7 Ring compression test – Prometal die382Die at room temp.Die at 450ºFDie at 450ºF (Polished)10 X With H-13 Dies showing recrystallization10 X With Prometal Dies showing friction and grain flow200 X With Prometal dies showing rapid quenchingFig. 8 Microstructures from ring compression samplesFig. 9 Scale Down Die Note –Fig. 10 Simulation for filling of scale down partThe above problem was modeled as axisymmetric and taking the dies as rigid and the material (Al-7075) at 800 F. The dies are at 450 F.REFERENCES1.

In this work an evaluation concept to analyze the thermal melt pool characteristics based on a high-speed camera is developed. The thermal radiation intensity of the melt pool is used to derive the thermal gradient and combined with an image rate of 41,000 fps the solidification rate is derived. The developed approach provides local data of the solidification for ML-based process adaptions but also serves for part individual quality assurance tasks.IntroductionAdditive Manufacturing (AM) has shown its potential of flexible manufacturing to replace cut off value chains during covid-pandemic since 2020 [2]. In addition, AM enables socially requested resource-efficient manufacturing of critical alloys, e.g. rare-earth-alloys [3]. During the past years metal-based AM has been established widely. When it comes to high quality with good reproducibility, PBF-LB is used [4].

rare-earth-alloys [3]. During the past years metal-based AM has been established widely. When it comes to high quality with good reproducibility, PBF-LB is used [4]. Based on a layer by layer applied powder bed, the part geometry is exposed using a laser beam to fuse the metal powder to a dense part. This melting process determines the resulting part properties and is therefore object of current research [5,6]. The layer-by-layer build up enables the use of process monitoring systems to observe and evaluate the process over the complete part. Furthermore, data driven approaches can be derived to control the process to achieve graded microstructure and therefore improved part functionality [6].Part properties can be estimated by its microstructure and geometry, including microstructural defects such as pores, cracks and chemical separation.

Most scientific works focus on process anomaly detection and therefore estimate defects whereas the evaluation of microstructural properties such as grain size is often neglected [7].The microstructural crystal growth depends on local cooling rate, solidification speed and nucleation [8]. The grow direction can be estimated by the temperature gradient (Figure 1). Concurring grain growth with misorientation leads to grain boundaries, which can be divided into low angle grain boundary (LAGB) and high angle grain boundary (HAGB) [9]. HAGBs have significant higher interfacial energy, which serves as resistor to dislocations and therefore increases1389mechanical strength.

HAGBs have significant higher interfacial energy, which serves as resistor to dislocations and therefore increases1389mechanical strength. Smaller grains lead to subsequently to more interfaces and therefore higher overall interfacial energy leading to higher mechanical strength [9].Gx R= konst.G/ R= konst.const.1,00E-041,00E-021,00E+001,00E+021,00E-011,00E+011,00E+03Rin mm/sGin K/mmconst.Figure 1: Graphical illustration of the correlation between solidification rate R, thermal gradient G and the resulting microstructure (based on [10])The solidification rate R and the temperature gradient G depend on the part geometry, material and process parameters (Figure 1). High volume energy input of 200 J/mm³ leads to bigger melt pools and therefore homogenous temperature gradients. This results in grains with an average length of 200 µm. Low energy input of 58 J/mm³ leads to smaller grains with a length up to 100 µm.

This results in grains with an average length of 200 µm. Low energy input of 58 J/mm³ leads to smaller grains with a length up to 100 µm. [11]To observe the process, multiple approaches have been developed [12,13]. Pyrometers (Melt Pool Monitoring, MPM) up to 100 kHz [14] are used to evaluate the process dynamics whereas thermographic cameras with up to 50 kHz [15] are used to evaluate the temperature. When it comes to multiphase-state where solid, liquid and vapor co-exist, an evaluation using a thermographic camera is difficult as it is only calibrated to a specified phase and temperature range [16]. Pyrometers are used to measure one value each time step, which does not represent the geometric thermal distribution of the melt pool. An evaluation of temperature gradients and cooling rates is therefore not possible. However, the recorded information are used to understand the process, e.g.

An evaluation of temperature gradients and cooling rates is therefore not possible. However, the recorded information are used to understand the process, e.g. melting behavior, and furthermore to adopt the process [17].1390State oftheart: Blob-DetectionCharacteristics:MeltpoolwidthMeltpoollengthNew approach: Phase detectionbasedon thermal intensitydistributionExtended characteristicsPixel in x-directionPixel in y-directionIntensityvsIsolineevaporationtemperatureT1IsolinesolidificationtemperatureT2SolidMeltKeyholewMPlMPSolidificationtemperatureT2Evaporation temperatureT1Figure 2: Schematic concept of the solidification evaluationFocus of this work is the evaluation of the solidification process using a high speed camera (HSC). The HSC is implemented on an PBF-LB manufacturing system and the influence of the process on the thermal radiation intensity evaluated.

The HSC is implemented on an PBF-LB manufacturing system and the influence of the process on the thermal radiation intensity evaluated. By defining two referenced intensity values, characteristics for the solidification rate and temperature gradient can be derived (Figure 2). Based on single melt tracks the method is validated.Research methodologyA commercially available PBF-LB/M system M290 from EOS GmbH (Krailling, Germany) is used to conduct the experiments. This system uses a 400 W single-mode fibre laser with a wavelength of 1060 nm focused to a spot size of 100 µm. The building volume has a size of 250 mm x 250 mm x 325 mm. Gas atomized stainless steel 316L powder is used for all experiments in this study. A scanning electron microscope (SEM) analysis of the powder is used to determine the particle distribution. Particles ranging from 11 to 54 μm and a d50 of 25 μm was determined.

A scanning electron microscope (SEM) analysis of the powder is used to determine the particle distribution. Particles ranging from 11 to 54 μm and a d50 of 25 μm was determined. The PBF-LB/M system is extended by an on-axis high speed camera system plasmoEye (plasmo Industrietechnik, Austria) as shown in Figure 3. The camera has a pixel size of 14 µm and a variable region of interest (ROI). The camera detects a wave length of 950 nm. The ROI is set to 80x40 pixel with a spatial resolution of 11.2 by 5.6 mm.

The camera has a pixel size of 14 µm and a variable region of interest (ROI). The camera detects a wave length of 950 nm. The ROI is set to 80x40 pixel with a spatial resolution of 11.2 by 5.6 mm. The recording rate is set to 41,000 fps and an external trigger signal from the melt pool monitoring system is used to time stamp each image.1391This time stamp is used to map each image with the position data from the melt pool monitoring system.MPM on axisFocusing unitData linkCameraKameraScanner unitBeam splitterMPMOn-Axis FotodiodeFokussiereinheitStrahlteilerDatenkabelScannereinheitFigure 3: Experimental Setup - On-Axis High-Speed-Camera mounted on an EOS M290The images are evaluated offline post process. Position- and exposure-dependent effects are taken into account by a correction and subsequent validation through a comparison of the generated melt track geometry with the evaluated melt track width.

The temporally and spatially resolved intensity distribution as well as the weld mode and the melt pool depth represent the input variables for the modeling of the process-microstructure correlation. In addition, the temperature gradient is determined via the spatial distance between the solidification and sublimation iso-lines and the knowledge of the melting and boiling temperatures. The solidification rate is determined via the change in position of the solidification front over time. The investigations and data recording start with the production of one-dimensional single melt tracks and are continuously extended to two-dimensional melt track composites and three-dimensional layer composites. The influence of the build position, the scan vector direction, and the process parameters on the detection of the melt track via defined isolines of the melt pool intensity is examined.

The influence of the build position, the scan vector direction, and the process parameters on the detection of the melt track via defined isolines of the melt pool intensity is examined. To evaluate the influence of these factors on the melt pool and the melt track detection, line tests are carried out and investigated by means of optical 3D measurement. The process parameters laser power and scan speed as well as the scan direction and the positioning in the build space are varied.The detectability of the melt pool using the blob detection algorithm strongly depends on the HSC noise. To reduce noise and therefore improve the quality of the blob detection, a noise correction is developed. The mean value of each pixel is evaluated over 1,000 images from different build jobs and subsequently the difference between reference mean value and current1392value derived for each image.

The mean value of each pixel is evaluated over 1,000 images from different build jobs and subsequently the difference between reference mean value and current1392value derived for each image. The amount of 1,000 images is set to get a proper database and is also usable for other materials.The basic algorithm for advanced evaluation is based on linking the melt pool geometry of the binary image with the intensity of the grayscale image (Figure 4). By overlaying the binary image with the intensity image, the noise is filtered out, as these would otherwise also be detected as meltpool. After superposition, the isolines can be formed for defined intensity values. This has the advantage that there is comparability across all experiments. The intensity level of the isolines can also be selected as a function of the exposure time in order to infer the intensity density, which enables a comparison of images taken at different exposure rates.

The isolines show a significant correlation to the different phase states present and were determined using a best fit method. Orthogonal to the direction of the first major axis, measurement lines are then drawn at regular intervals along the length of the melt pool and the intersection points between the measurement line and the isoline are determined.

The distance between the intersections on a measuring line is defined as the melt track width of the measuring point.IntensityPixel in x-directionPixel in y-directionPixel in y-directionGrayscaleimageMelt pool detectionIntensityisolinesMelt pool measurementsCharacteristicsRaw imageHSCPixel in y-directionPixel in x-directionPixel in x-directionIntensityFigure 4: Schematic images of the central evaluation stepsThe evaluation of the cooling rate via a comparative quantity from the HSC data is defined via the isoline of the keyhole edge (if a keyhole is present) and the solidification line (isoline of the intensity I = 0.2 digits), which is of decisive relevance for the geometry. The temperature gradient corresponds to the difference between the melting temperature and the evaporation temperature and depends on the distance between the melting transition and the evaporation transition. For this purpose, a direction-dependent variable is defined.

For this purpose, a direction-dependent variable is defined. The cooling rate is defined by the time difference between the transition of the gaseous state at the keyhole edge to the melt and the solidification of the melt and is to be determined locally.1393For the metrological calibration, three melt tracks are produced on a cuboid. The edge length of the cuboid is 10 mm. The distance between adjacent melt tracks is 3 mm, the distance to the edge of the cuboid is 2 mm. One side of the cuboid is parallel to each of the melt tracks.x-Position in µmHeight in µmTrack centerliney-Position in µmParameter set1, scanangle 135°Referencetrack centerlineFigure 5: Example of the optical melt track measurement referenced by the intersection of the melt track and the reference trackFor metrological referencing of the origin of the melt tracks, a reference melt track was manufactured perpendicular to the melt tracks at a distance of 0.5 mm from the edge of the cuboid (Figure 5).

This thus crosses the melt tracks to be measured and can therefore be used as a starting point for measurement purposes in the height profile and by locating the process monitoring data. To evaluate the dependence of position and scan direction, eight cuboids per position were produced with varying scan directions at nine positions. The scan direction varies between 0° and 315° with an increment of 45°. The position varies evenly distributed in nine fields, three each in x-direction and three in y-direction. The basic influence of the process parameters on the detection and calibration of the isolines is done by varying the laser power and the scanning speed. The process parameter set qualified by the system manufacturer with a line energy of 0.18 J/mm is used as reference parameter set.

The process parameter set qualified by the system manufacturer with a line energy of 0.18 J/mm is used as reference parameter set. A process parameter combination with reduced line energy of 0.09 J/mm in the range of heat conduction welding and a process parameter combination with increased line energy of 0.36 J/mm in the range of keyhole welding are selected.ResultsThe calculated mean noise of each pixel over 1,000 images of the HSC and subsequent consideration is used to decrease the noise by 85 % to 90 %. While isolines larger than the intensity of 0.4 digits can be interpreted as gaseous phase, isolines of intensities between 0.2 and 0.4 digits indicate melt. Isolines of intensities smaller than 0.2 digits indicate partially or completely solidified melt. Characteristic of the gaseous phase is the high fluctuation of the cross-sectional shape. An isoline stabilized in the cross-sectional shape near the keyhole indicates the transition between melt and keyhole.

An isoline stabilized in the cross-sectional shape near the keyhole indicates the transition between melt and keyhole. The transition between melt and solidified material can be seen by the transition into noise. From these isolines, further characteristic parameters are defined. These are the major axis lengths of the individual isolines as well as the width along the first major axis at defined measurement intervals. However, not all isolines are equally suitable for determining the melt pool width at defined measuring intervals. In particular, the isolines of intensity 0.1 to13940.2 digits exhibit a large dependence of the noise behavior and corresponding irregularities.

The measuring points of parameter set 1 as well as the standard deviation overlap with the metrologically determined standard deviation. For all parameter sets, a large scattering of the measurement points can be seen.y-PositionY0Y-Parameter set1x-PositionY+X0X-PositionX+ShieldinggasMelt track widthin µmMean widthHSCMeasurement pointswidthHSCMean widthAliconaStandard deviationAlicona1395Figure 7: Comparison of melt track width for different positions based on 3D-optical measurement (off-line, marked in red) and the high speed camera approach (in-line, marked in grey)The observation of the melt track detection via the isolines of the melt pool geometry shows a dependence on the position. The average melt track width is on average lower than the measured melt track width (Figure 7). Melt tracks which were produced on the X- position are detected as smaller by optical process monitoring compared to melt tracks which were produced on the X+ positions.

Melt tracks which were produced on the X- position are detected as smaller by optical process monitoring compared to melt tracks which were produced on the X+ positions. A dependence on the y-position is not statistically significant.The mean intensity shows high position dependency and varies between 0.03 and 0.07 digits. The X-Y- position shows the lowest mean intensity (0.03-0.04 digits). X-Y+ and X+ show the highest mean intensity (0.055-0.07 digits) as shown in Figure 8. The position-dependent intensity is used to adopt the characteristic intensity values for the evaluation. The characteristic intensity value for the transition from liquid to solid is set to 90% of the position depended mean intensity.

The characteristic intensity value for the transition from liquid to solid is set to 90% of the position depended mean intensity. The intensity value for the transition from liquid to gaseous is set to 90% of the position dependent maximum intensity.x-Position in pixelMean intensityxData pointsy-Position in pixelFigure 8: Influence of the position of the build platform on the detected mean intensityA higher influence of the scan angle is seen when higher line energy density is used. The influence on the melt track evaluation is mainly for 90° (laser moves against shielding gas direction) and between 235° and 330° (laser is moving in shielding gas direction). The detected melt track via optical monitoring increases compared to other scan angles (Figure 9).

The detected melt track via optical monitoring increases compared to other scan angles (Figure 9). The influence of the scan angle depends on the parameter set.1396Mean widthHSCMeasurement pointswidthHSCMean widthAliconaStandard deviationAliconaScan vectorangle in °Parameter set1Melt track widthin µmFigure 9: Comparison of melt track width for different scan angles based on 3D-optical measurement (off-line, marked in red) and the high speed camera approach (marked in grey)Finally, the corrections based on position-dependency, scan vector angle, and noise were used to adopt the melt pool evaluation. This results in improved characteristics to characterize the melt pool and the resulting solidification.

This results in improved characteristics to characterize the melt pool and the resulting solidification. Deviations in the melt track are also detectable using the new approach (Figure 10).Parameter set1, scanangle 135°y-Position in µmHeight in µmx-Position in µmxMelt track widthHSCFigure 10: Illustration of the melt track width determination based on the HSC (red crosses) compared to the height profile of the melt trackOverall, the accuracy was improved, and the influence of position and scan angle was reduced (Figure 11). The deviation between melt track and evaluated melt pool characteristics was reduced to 1-5 µm. There are two outliers (180° and 225°) using parameter set 1 with a deviation higher 5 µm. However, the influence of the parameter set still exists. The evaluated melt pool1397characteristics show higher results than the real melt track width.

However, the influence of the parameter set still exists. The evaluated melt pool1397characteristics show higher results than the real melt track width. The deviation is between 4 to 20 µm.Melt track widthAliconaMelt track widthHSCMelt track widthin µmParameter setScan vectorangleFigure 11: Final results of the melt track evaluation based on the HSC (blue crosses) compared to the optical determined melt track width for different parameter sets and scan anglesDiscussionThe detection of the melt pool geometry is based on the intensity distribution. The emission coefficient of the emitted thermal radiation of the melt pool depends on the material, temperature, and phase. Assuming homogeneous energy coupling, a uniform melt pool can be assumed to emit similar thermal radiation. The intensity of the detected thermal radiation would be constant under this assumption.

Assuming homogeneous energy coupling, a uniform melt pool can be assumed to emit similar thermal radiation. The intensity of the detected thermal radiation would be constant under this assumption. This assumption neglects influencing factors of the energy coupling as well as changed detectability of the thermal radiation. The energy coupling is reduced by the angle of incidence of the laser beam and the scattering of the laser beam due to the formation of smoke above the vapor capillary and thus causes variable melt pool characteristics. In addition, the smoke causes scattering of the emitted heat radiation and thus influences the detectable intensity of the thermal radiation.The effects investigated must be taken into account in the evaluation. For this, a separation of the effects must take place and a correction or adjustment of the evaluation algorithm must be made. Both the energy coupling and the heat radiation are position dependent.

For this, a separation of the effects must take place and a correction or adjustment of the evaluation algorithm must be made. Both the energy coupling and the heat radiation are position dependent. The consideration of the resulting fumes is dependent on both the scan angle of the laser beam and the process parameters. In the process, the stripe angle or the scan vector angle influences the interaction between the laser beam and the generated fume. If the laser travels in the opposite direction to the shielding gas direction, the fume generated in the interaction zone is transported out of the following interaction zone by the shielding gases. In the opposite case, when the scan vector direction and the shielding gas direction are identical, the fume is transported by the shielding gas into the new interaction zone and changes the laser beam properties geometry and power by scattering and reflection.

Thus, deviating material properties can result from the changed laser- material interaction. The extent of the change depends on the combination of process parameters1398and the inert gas flow. Furthermore, not only the laser radiation is changed, but also the detectable radiation in the process monitoring systems. Similar to the laser beam, the radiation emitted by the melt pool is modified by the fume. Scattering and reflection occur, with the result that the process radiation is less perceptible. All other alignments between the travel path of the laser beam and the shielding gas flow direction can be described by a partial scattering of the laser beam. The process parameters influence the quantity of the resulting fumes and thus the degree of scattering. High line energies tend to lead to greater fume generation due to vaporization of alloying elements.

High line energies tend to lead to greater fume generation due to vaporization of alloying elements. The scattering of the emitted heat radiation of the molten bath by flue gases is equally dependent on these same process phenomena.The deviation of the melt track width for process parameter set 2 can be explained by the increased scanning speed. On the one hand, this results in a shorter exposure time and thus a longer melt pool length, and on the other hand, a higher scanning speed leads to faster removal from the area of welding fumes, which influence the laser radiation and the heat radiation. This prevents scattering of the laser radiation and the heat radiation by fumes.Summary and outlookIn this work a new approach to determine local thermal solidification characteristics is presented. A high speed camera with 41,000 fps is used to detect thermal radiation of the process zone in the wave length of 950 nm.

A high speed camera with 41,000 fps is used to detect thermal radiation of the process zone in the wave length of 950 nm. The evaluation of the images is based on a melt pool detection and identification of intensity isolines, which define the characteristic transitions from gaseous to liquid and from liquid to solid phase. The experimental validation uses a best fit to compare metrological measurements with the derived melt track width from the HSC. The position, scan angle and process parameters are considered using correction factors within the evaluation. This results in good predicting quality of the melt track width with a low deviation less 5 µm except for two outliers. To enhance the model quality and identify the limits of the method, more experimental investigations have to be made.For further research the extension of the evaluation algorithm is focused to derive the the thermal gradient. This concept has to be validated experimentally.

This concept has to be validated experimentally. Therefore, specimens have to be evaluated in terms of microstructural characteristics, e.g. by means of Electron-Backscatter- Microscopy. The current two-dimensional evaluation is supposed to include the melt track depth. Therefore, a model to predict the melt pool depth can be derived based on metallographic investigations or already established models. The ability to transfer to further materials has to be investigated. Based on the proposed approach, the experimental setup as well as the experimental design and the measurements can be copied to new systems and materials.The overall aim of enabling data-driven processes is now provided with an approach to generate a high quality database. The sampling rate is able to match the process speed. The amount of information covers geometrical aspects and information which can be controlled by process adaptions.

The sampling rate is able to match the process speed. The amount of information covers geometrical aspects and information which can be controlled by process adaptions. The developed approach provides local data of the solidification for ML-based process adaptions but also serves for part individual quality assurance tasks1399AcknowledgmentsThis research and development project is funded by the Federal Ministry for Economics and Climate Action (BMBK) within the “Technologietransfer-Programm Leichtbau” Program (funding number 03LB2009C) and implemented by the Project Management Agency Jülich (PTKA). The author is responsible for the content of this publication.References[1] DIN EN ISO/ASTM 52900:2022-03, Additive Fertigung_- Grundlagen_- Terminologie(ISO/ASTM 52900:2021); Deutsche Fassung EN_ISO/ASTM 52900:2021, Beuth Verlag GmbH, Berlin.[2] R.C. Advincula, J.R.C. Dizon, Q. Chen, I. Niu, J. Chung, L. Kilpatrick, R.

viewed PaperEvaluation of the Ecological Footprint for Parts from AlSi10Mg manufactured by Laser Powder Bed FusionC. Weiss*1, C. Boedger2, E. Schiefer3, D. Heussen1, C.L. Haefner1,41Fraunhofer-Institute for Laser Technology ILT, Germany, 52074 Aachen 2Paderborn University, Automotive Lightweight Design, Germany, 33100 Paderborn 3Frankfurt University of Applied Sciences, Product development and EcoDesign, Germany, 60318 Frankfurt 4RWTH Aachen University, Chair for Laser Technology, Germany, 52074 AachenCorresponding author: christian.weiss@ilt.fraunhofer.deAbstractThe manufacturing industry contributes immensely to the global emissions and therefore is a key factor that has to be addressed when a more sustainable production is desired. Laser Powder Bed Fusion (LPBF) is an AM technique that offers the possibility to manufacture metal parts in a more material efficient way due to the layer-by-layer build-up.

Laser Powder Bed Fusion (LPBF) is an AM technique that offers the possibility to manufacture metal parts in a more material efficient way due to the layer-by-layer build-up. Nevertheless, the processing chain for parts from LPBF contains additional steps like powder atomization, which also influence the ecological footprint of the production chain. Within this work, a life-cycle model for the production step of parts from AlSi10Mg powder material is developed. The model is supplied with data from the powder atomization up to the production step, either by literature, database or experimental measurements during production. The footprint in terms of CO2 emissions is then analyzed and emission-intense steps are identified. Two manufacturing scenarios are considered to evaluate the sensitivity on the emissions.IntroductionOne of the major challenges of recent times is dealing with the man-made climate change.

Two manufacturing scenarios are considered to evaluate the sensitivity on the emissions.IntroductionOne of the major challenges of recent times is dealing with the man-made climate change. The main driver of climate change is based on CO2 (carbon dioxide) emissions [1]. As part of the "Agenda2030," the United Nations (UN) has defined 17 goals for sustainable development, one of which includes sustainable manufacturing of products [2]. In Germany, for example, industry is responsible for approximately 30% of the country's total energy consumption and associated emissions [3]. Despite efforts by the German government to reduce emissions, the level for the industrial sector has remained almost constant since 2009 until today [4]. This puts the German government's stated goal of becoming largely climate-neutral by 2050 at great risk [5].

This puts the German government's stated goal of becoming largely climate-neutral by 2050 at great risk [5]. Therefore, there is a need for action, among others in the industrial sector, to counteract this trend and to support the achievement of the climate targets in a sustainable manner.Lightweight design and manufacturing of components can contribute to the conservation of resources, raw materials and energy. The weight- and performance-optimized design of components can reduce climate-damaging emissions both during production and during the use of the components. In addition to the ecological benefits, lightweight design also offers economic potential [6]. Additive manufacturing (AM) processes in particular offer the possibility to produce geometrically optimized lightweight structures. Function-integrating and topologically optimized921high-performance metal components are primarily manufactured using the Laser Powder Bed Fusion (LPBF) process.

Function-integrating and topologically optimized921high-performance metal components are primarily manufactured using the Laser Powder Bed Fusion (LPBF) process. In this manufacturing process, powdered raw material is melted locally by means of a laser beam, thus producing a solid body. Unmelted powder material can be reused after the process through reconditioning procedures. The industry has recognized the great potential of this process and is already applying the manufacturing technology in a wide field of sectors such as aerospace and automotive. At the current time, though, a lot of work is going into the further development of the additive manufacturing technology to increase productivity as well as to expand the materials that can be processed in LPBF [7].Under the current challenge of climate protection, the manufacturing industry plays an important role.

A holistic view of the emissions of process chains from the raw material to the end of the life cycle (end-of-life) is therefore necessary in terms of sustainability. However, the LPBF method has so far mostly only been quantified in terms of economic benefits [8]. The consideration of sustainability or ecology with regard to energy and resource saving potentials, on the other hand, has so far only been of secondary relevance. Life cycle assessments (LCA) of process chains can help to generate a deeper understanding of the influence of the sustainability of LPBF- manufactured components through holistic balancing. However, since existing solutions for life cycle analysis only partially consider the LPBF process chain (usually only on the basis of the manufacturing process), there is a need for action to fill this knowledge gap.

Furthermore, the methodology of LCA is usually limited to products and has not yet been used as standard in manufacturing [9].Most available publications on LCA of LPBF focus mainly on the manufacturing stage of the respective components. It was consistently found that the electrical energy consumed is the largest factor with regards to sustainability in this phase. As an example, Huang et al. [10] analyze the influence of the different materials on energy consumption. Due to the lower melting point of aluminum, 73 - 94 MJ/kg are consumed to produce a component. In comparison, 110 - 141 MJ/kg of electrical energy must be expended for a titanium alloy. The publications also show that the figures for electricity consumption are highly dependent on the used system.

The publications also show that the figures for electricity consumption are highly dependent on the used system. On another LPBF system (Concept Laser M3 Linear), it can be seen that the consumptions are not only higher, but exceed them by a factor of five (aluminum: 374 - 520 MJ/kg and titanium: 561 - 780 MJ/kg).Neither found publication considers the maintenance and cleaning of the equipment used. Furthermore, the utilization phase of the components is hardly considered either, since they are mostly individual parts that only form a holistic product after assembly. Böcking and Tillman [11] take an in-depth look at utilization and found that AM brings potential benefits for the environment and resources, but also disadvantages that need to be circumvented. In the application example of a Volvo diesel engine, the reduced impacts in the use phase more than offset the increased impacts in production.

In the application example of a Volvo diesel engine, the reduced impacts in the use phase more than offset the increased impacts in production. This is due to lightweight-optimized components, which reduces fuel consumption and thus lifecycle impacts. Similar effects are observed for all products with high and weight- dependent energy consumption in the use phase. According to the authors, the results are comparable with similar assessments in the literature.Since investigations along the rather complex LPBF process chain are scarce, the present work will first establish the product life cycle of a component manufactured using the LPBF technology.

Based on databases, publications and experimental measurements, the LCA will then922be established to holistically consider and analyze the environmental impacts and thus identify the ecological footprint of this AM technique.Introduction to Life-Cycle AssessmentLife Cycle Assessment is a method used to assess the potential environmental impact of the production and use of products. LCA thus considers the entire life cycle of the product, from raw material extraction to disposal. Due to increasing awareness for environmental sustainability, this method has been developed and is gaining more and more importance [12]. Life cycle assessment can help to find ways to improve the environmental characteristics of the product and thus create information for decision making.

Life cycle assessment can help to find ways to improve the environmental characteristics of the product and thus create information for decision making. Aditionally, life cycle assessment also serves marketing through the implementation of environmental labels [13].When analyzing a product life cycle, a distinction is made between mainly three differentapproaches with varying extents [13]:Cradle-to-grave analysis includes the complete life cycle assessment from resource extraction to final disposalCradle-to-gate is an assessment of a specific product life cycle, which is defined from resource extraction to product manufacturingGate-to-gate, which is a partial LCA that considers only one value-adding process in the entire production chain.In general, the LCA analysis can be divided into four phases, which are closesly connected with each other. Figure 1 shows these phases and interrelationships.

Figure 1 shows these phases and interrelationships. These are explained in more detail below.Figure 1: The four phases of Life Cycle Assessment923Phase 1: Definition of the objective and the scope of the investigationIn the first phase, the system boundaries and the functional unit are defined. The latter is hereby defined as the " quantified performance of a product system for use as a reference unit" and determines what good is being investigated [13]. The subsequent analyses then relate to this functional unit. If individual process steps or even life cycle stages of the product are not included in the scope of the assessment, this must be justified in this stage as well [13]. In addition, the necessity of the study is justified in this phase [14]. Likewise, the various system boundaries, which can be divided into technical, geographical and temporal boundaries, must be specified in this step.

Likewise, the various system boundaries, which can be divided into technical, geographical and temporal boundaries, must be specified in this step. Depending on the data availability, possible cut-off criteria and simplification must then be applied [15].Phase 2: Life cycle inventory (LCI)In the life cycle inventory, the inputs and outputs of the system under investigation are determined. This results in a collection of the data necessary to achieve the objective of the investigation. If necessary, these data must be standardized to the previously defined functional unit by calculation procedures [13].The input variables (raw materials, auxiliary materials, etc.) are often related to the output variables (pollutants, waste, etc.). The outputs can be further differentiated into "desired" and "undesired".

The outputs can be further differentiated into "desired" and "undesired". “Desired” outputs are mostly semi-finished products, components or entire products, while “undesired” outputs are almost unavoidable like emissions and waste [16].The data that is entered into a LCI can be divided into two different data types: generic data and primary data. Generic data is composed of available data sets from material and process databases, which are integrated in a computer-based tool. Primary data are collected experimentally for process steps. It is important at this point that the data sets are sufficiently documented and thus comprehensible to third parties [15]. Based on the collected data of the different processes, a product system can subsequently be created. Ideally, this is modeled in such a way that the inputs and outputs at the boundaries are elementary flows. Elementary flows are the substances or energies that are added to or leave the defined system.

Elementary flows are the substances or energies that are added to or leave the defined system. Due to the complex products and the resulting strong interconnectedness of the systems, this is rarely possible [13]. In order to present such a product system in a more manageable way, these are usually structured according to technical subsystems, which represent successive process modules.Phase 3: Impact assessmentIn the third phase, the Life Cycle Impact Assessment (LCIA), the previously generated data is examined and evaluated with regards to environmental impacts. After the selection of impact categories and impact indicators (quantifiable representation of an impact category), the most important elements of this phase include classification and characterization [13]. In the classification, resource consumption and emissions are divided into categories according to their effects on the environment, such as greenhouse effect, human toxicity or eutrophication.

In the classification, resource consumption and emissions are divided into categories according to their effects on the environment, such as greenhouse effect, human toxicity or eutrophication. It is also possible that an emission has several effects in different impact categories. The characterization takes place within an impact category. Here, the emissions and consumptions are weighted with a924so-called equivalence factor, which are related to a reference substance in order to make them comparable [14].Phase 4: EvaluationEvaluation is the final phase in which the previously obtained results are summarized in accordance with the objective as a basis for conclusions or recommendations. One of the main objectives is always to identify main influencing factors [17].Goal and Scope DefinitionThe aim of this work is to analyze the influences of the input and output variables generated by components made of AlSi10Mg, manufactured by the LPBF process.

The effects of the different product life cycle stages from raw material extraction to manufacturing (cradle-to-gate) are considered. Within the life cycle phases and stages described, intermediate steps such as raw material and powder production, transportation and others are considered in addition to the technical factors associated with each operation. Data is collected in a number of ways. Data is either collected experimentally the process modules or, if this is not possible, supplemented with extensive literature research and databases in the software used (OpenLCA). The determined mass values along the entire process chain are measured with a scale (measuring accuracy 0.1 kg). The meter used to record the power consumption is a Fluke 435 series II, which is connected between the equipment and the power supply.The investigated component is an impeller. The diameter of the component is 100 mm with a height of 35 mm, a component volume of 44.3 mm³ and a weight of 119 g.

The diameter of the component is 100 mm with a height of 35 mm, a component volume of 44.3 mm³ and a weight of 119 g. The CAD data for the impeller is openly available at GRABCAD [18] and rescaled by a factor of 0.5.For the definition of the functional units, two scenarios are considered in order toinvestigate influences of packing density during LPBF, which are represented in Figure 2.:1. the production with the LPBF process of one impeller per build job from AlSi10Mg, assuming that production takes place once per working day over 8 years.2. the production with the LPBF process of four impellers per build job from AlSi10Mg, under the assumption that production takes place once per working day over 8 years.In this work, the cradle-to-gate scope ranges throughout the individual steps that are further described in the following and depicted in Figure 3. The beginning of this life cycle assessment is the extraction of the raw materials aluminum, silicon, magnesium, etc.

The beginning of this life cycle assessment is the extraction of the raw materials aluminum, silicon, magnesium, etc. and their transportation. Based on this, the production of the AlSi10Mg alloy is considered. Since the LPBF process requires the material to be in powder form, it is essential to also consider the powder production process by means of gas atomization. In the LPBF process, the defined component is manufactured from the powder material and data is recorded. In the present work, , the manufacturing of the LPBF system itself and its cleaning after processing are also considered. After the LBPF process, the powder that has not been melted is cleaned of process-induced by-products, such as spatter, using a sieving system and returned to the process chain. The build job is then heat treated to increase strength and reduce residual stresses. To complete the manufacturing step, the components are separated from the build platform using a milling machine.

The build job is then heat treated to increase strength and reduce residual stresses. To complete the manufacturing step, the components are separated from the build platform using a milling machine. The subsequent use phase and disposal of the components are not accounted for. This results in the examination frame (red) in Figure 3.925a)b)Figure 2. Illustration of investigated scenarios for experimental data generation; a) single build; b) maximum build (SG = shielding gas, REC = recoating direction, BD = build direction)The process steps mentioned are often connected with upstream chains. This means that the auxiliary materials and energy flows used are also considered in the balancing. Thus, the two life cycle assessments present all relevant material and energy flows out of and into the environment, with reference to the defined functional unit.Figure 3.

Thus, the two life cycle assessments present all relevant material and energy flows out of and into the environment, with reference to the defined functional unit.Figure 3. System boundary of the LCA study along the typical LPBF process chain926Assumptions for the investigations:Due to the large number of processes and associated raw materials, semi-finished products, etc. on the one hand, and limited time and human resources on the other, it is unavoidable to make assumptions and constraints for the product system. Without assumptions, many process steps would be omitted and thus it would not be possible to estimate whether the resulting pollutant emissions are relevant [12]. In general, it must also be mentioned that the data sets used from the ecoinvent database are also based on case-specific framework conditions.

In general, it must also be mentioned that the data sets used from the ecoinvent database are also based on case-specific framework conditions. A detailed explanation of the assumptions is given during the preparation of the life cycle inventory in the respective process step in the LCI.Life Cycle Inventory (LCI) for LPBF Process ChainRaw Material Extraction & Alloying:The individual weight-percentage [wt.-%] for composition of the alloy must first be extracted, processed and melted to produce the AlSi10Mg raw material. For this process, no information can be found in the literature regarding the electrical energy consumed. In the ecoinvent database used, there is only one product system for the production of AlMg3 alloy. Therefore, a simplification is made at this point. The AlMg3 data set is modified to the raw materials of the AlSi10Mg alloy.From the used database, a value of 13.86 kWh per kilogram for the extraction of primary aluminum (Al) can be found.

The AlMg3 data set is modified to the raw materials of the AlSi10Mg alloy.From the used database, a value of 13.86 kWh per kilogram for the extraction of primary aluminum (Al) can be found. For the same quantity of silicon (Si) and magnesium (Mg), 11 kWh/kg and 1.52 kWh/kg are used, respectively. Under consideration of the wt.-% for the investigated aluminum alloy, the total consumption for the AlSi10Mg alloy is 13.52 kWh/kg. Comparatively, 5.2 kWh/kg are consumed for the production of steel [19] . After this process step, the AlSi10Mg alloy is available as a semi-finished product.For scenario 1, 44 kg of the alloy are required to produce 13.2 kg of powder material with a particle size of 20 – 63 μm. This results in an electricity consumption of 594.83 kWh and is accounted for with the Chinese electricity mix. In the second scenario, not one but four impellers are manufactured additively in one build job. This results in different inputs and outputs in the life cycle inventory.

In the second scenario, not one but four impellers are manufactured additively in one build job. This results in different inputs and outputs in the life cycle inventory. The raw material extraction and alloy production entails an electrical energy consumption of 599.29 kWh to produce the required quantity of 44.33 kg AlSi10Mg raw material. Equivalently, the inputs of the individual raw materials also increase.Powder Atomization:For powder production by gas atomization, in addition to the raw material, the electrical energy and nitrogen consumption must be declared as inputs. Since atomization is carried out externally, and manufacturers do not publish precise details, available literature studies are used to collect data. Aluminum alloys mostly are atomized via gas atomization due to the high reactivity of the alloy in powder state [20].927Faludi et al. [21] consider the production of an aluminum alloy using gas atomization.

[21] consider the production of an aluminum alloy using gas atomization. Here, the electrical energy requirement is put at 8.1 MJ/kg (2.25 kWh/kg). For this and the following process steps, the European electricity mix is used. Nitrogen is used as atomization medium in this specific application. According to [22], a liquid metal flow of 20 kg/min consumes approx. 8 m³ of gas per minute. From this, a gas volume of 0.4 m³ per kg can be derived.Since it is not possible to gain complete control over the particle size of the atomized powder, the distribution can be influenced by varying the ratio of gas to melt flow rate. Research in the field of gas atomization shows that finer particle size distributions can be achieved by using hot gas atomization [20]. The required particle size of 20 - 63 μm accounts for only about 30 % of the total particle size distribution [23] (Figure 4 b). The remaining 70 % are used for other processes such as isostatic pressing.

The remaining 70 % are used for other processes such as isostatic pressing. Thus, for one kilogram of AlSi10Mg powder in the required particle size, more than three kilograms of raw material have to be atomized. The nitrogen gas, which is relevant for the atomization process, is purchased from a nearby company located in Austria. For this purpose, an assumption of 100 km is accounted for in the process, which is covered by a truck. The AlSi10Mg powder produced at the end of the process is transported by truck over a distance of approx. 970 km to the production site in Aachen, Germany.The raw material, which is atomized with 22.16 kg nitrogen, requires 99.75 kWh for the production of 13.3 kg AlSi10Mg powder, with a particle size of 20 - 63 μm.

The remaining 70% of the powder (> 63 μm) will not be considered further in this scope of investigation, as it can be used for other manufacturing processes.a)b)Chemical composition [wt.-%] Si Mg Al 0,34 9,6 Balance Mn Cu Ni >0,01 >0,01 >0,01 >0,01 >0,01Fe 0,11 PbZn 0,02 SnPropertiesParticle size Atomization medium20-63 µm NitrogenFigure 4. a) Chemical composition and properties of AlSi10Mg powder for this study; b) typical yield of gas powder atomization and LPBF range (marked yellow), adapted from [24]LPBF machine & LPBF manufacturing:It is not usually possible to account for the manufacturing of a complex machine, such as an LPBF system in this case. One reason for this is that the exact manufacturing steps and materials used represent internal company information that is protected accordingly.

One reason for this is that the exact manufacturing steps and materials used represent internal company information that is protected accordingly. One option would be to disassemble the available plant itself and thus estimate the production steps with the materials.928However, the effort and the risk of not reassembling the system in a functional way are high. However, since manufacturing of the system should not be neglected in the consideration of the entire process chain, literature and data thanks to Faludi et al. [21] is used. As the Renishaw AM250 (investigated by Faludi et al. [21]) and the SLM Solutions 280HL (used in this work) have a similar overall size and weight, the data from the Renishaw in openLCA is transferred to the present work. At this point, therefore, a simplification of the analysis takes place, since the assumption is made that the two machines have an identical design.

At this point, therefore, a simplification of the analysis takes place, since the assumption is made that the two machines have an identical design. Only the laser components were considered twice in the balancing.Since the LPBF manufacturing process is a complex one, a large number of inputs and outputs have to be considered. In addition to the equipment and CAD component already mentioned, other inputs must be considered. The aluminum building platform has a weight of 8.3 kg. Before starting the build job, the build area must be flooded with the inert gas (argon). Approximately 3.52 kg or 1700 liters are required for this [25]. A standard parameter from the machine manufacturer is used for manufacturing. During manufacturing, the overpressure in the build chamber of 30 mbar is slightly readjusted by the machine itself, since the system is not 100% sealed.

During manufacturing, the overpressure in the build chamber of 30 mbar is slightly readjusted by the machine itself, since the system is not 100% sealed. Based on previous scientific work, this is calculated to be 1.357 g/min.An overview for the needed durations and measured power consumptions for both investigated scenarios is given in Table 1. A more detailed overview of the inputs and outputs of the LPBF production step is shown in Table 2. The 51.16 kWh of consumed electrical energy can be divided. During the preparation time of 1 h 23 min, the build platform is aligned and preheated to TPre=200 °C. In addition, the build area is flooded with argon during this time. This results in a consumption of 3.78 kWh. The build process takes 4 h 34 min from the first to the last exposure, requiring 16.41 kWh (exposure time 2 h 9 min, coating time 2 h 24 min from machine log files). The argon consumption amounts to 3.52 kg.

The argon consumption amounts to 3.52 kg. In the subsequent post-processing time of one hour, the build platform cools down and the powder that has not melted is removed. During the downtime, the system is not disconnected from the power supply, consuming 1.74 kWh per hour. The high downtime results from the assumption that only one build job is produced per day. This time results in a power consumption of 29.24 kWh. The output of the build job, consisting of the component and the construction platform, is 8.5 kg. The unused powder, which is present in the overflow container at the end of the process (12.7 kg), is carried on for reprocessing. The argon evaporates as soon as the build chamber is opened.For scenario 2, many inputs of the LPBF process remain unchanged, such as the required equipment and transport routes. Of great interest in this case is the amount of electrical energy consumed, which is measured at 59.35 kWh.

Of great interest in this case is the amount of electrical energy consumed, which is measured at 59.35 kWh. In the preparation time of 1 h 41 min, which serves for the alignment and preheating of the construction platform as well as the flooding of the build chamber with argon, there is a consumption of 5.076 kWh. The build process takes 6 h 43 min and requires 27.177 kWh (exposure time 4 h 16 min, coating time 2 h 23 min) to simultaneously produce four components with two lasers. The post-processing of the manufacturing process takes one hour. During the subsequent downtime of 14 h 36 min, the system is not taken off power, consuming 1.742 kWh per hour. The argon consumption amounts to 3.798 kg. In this scenario, as well as before, the used silicone lip in the recoater has to be disposed of at the end due to wear. On the output side, the weight of the build job (build platform + components) is 8.9 kg. The non- solidified powder residues have a weight of 12.4 kg.

On the output side, the weight of the build job (build platform + components) is 8.9 kg. The non- solidified powder residues have a weight of 12.4 kg. As in the first scenario, the argon929evaporates as soon as the build chamber is opened.Scenario 1Scenario 2Preparation LPBF processing Post-processing Idle TotalDuration Power consumption Duration 01 h 41 min 03.78 kWh 01 h 23 min 06 h 43 min 16.41 kWh 04 h 34 min 01 h 00 min 01.73 kWh 01 h 00 min 14 h 36 min 29.24 kWh 17 h 03 min 24 h 00 min 51.16 kWh 24 h 00 minPower consumption 05.07 kWh 27.17 kWh 01.73 kWh 25.26 kWh 59.26 kWhTable 1. Time and power consumption overview for scenario 1 and scenario 2Heat treatment:The heat treatment is carried out according to the T6 process. In this process, the construction job is first heated to 525 °C for five hours, cooled in water and then aged at a temperature of 165 °C for seven hours. According to Scharf et al.

In this process, the construction job is first heated to 525 °C for five hours, cooled in water and then aged at a temperature of 165 °C for seven hours. According to Scharf et al. [26], this step in the process chain requires 25.96 kWh of electrical energy. The cooling basin has a volume of 50 liters. The evaporation of the water during the interim quenching is accounted for with one kilogram per construction job.

The evaporation of the water during the interim quenching is accounted for with one kilogram per construction job. After treatment, the build job is taken to a nearby company for further processing.LPBF manufacturingInputsOutputsDescriptionAmountSourceDescriptionAmountSourceScenario 1Scenario 2Scenario 1 Scenario 2LPBF system+1+1Faludi [21]Build Job8.5 kg8.9 kgExp.LPBF cooling system+1+1Faludi [21]Powder remains12.7 kg12.4 kgExp.CAD part+52.05 cm³+208.21 cm³Materialise MagicsArgon3.52 kg3.79 kgExp.Aluminum substrate+8.3 kg+8.3 kgExp.Silicone lip0.028 kg0.028 kgExp.AlSi10Mg powder+13.20 kg+13.30 kgExp.Electrical power+51.16 kWh +59.26 kWhExp.Argon+3.52 kg+3.79 kgExp.Silicone lip+0.028 kg+0.028 kgExp.Table 2: Detailed overview of input and output data for the LPBF manufacturing steps in scenario 1 and scenario 2 (Exp.

= Experimentally measured)Machining930The balancing of the post-processing is carried out exclusively with the database and without the electrical discharge machining (EDM) procedure. The amounts of electrical energy (0.336 kWh) and the coolant are deposited depending on the material waste. The material waste here refers to the weight of the support structure to be removed of 63.31 g (scenario 1). After a build job is defined as an input, this results in two outputs, firstly the component without the support structure and secondly the construction platform, which can then be reused.Reworking and separating the impellers in scenario 2 from the build platform requires 1.22 kWh of electrical energy, as well as some coolant. Wear of indexable inserts is not considered, as the AlSi10Mg alloy has quite soft material properties. Outputs are defined as the four components (0.476 kg) and the build platform.

Wear of indexable inserts is not considered, as the AlSi10Mg alloy has quite soft material properties. Outputs are defined as the four components (0.476 kg) and the build platform. The material waste is composed of the removed support structures and the face milling of the build platform.Impact Assessment and InterpretationTo perform an impact assessment (LCIA) with openLCA, the impact assessment method must be selected at the beginning. In this case, the CML baseline method is used. The basis of the calculation is the classification and the characterization of resource consumptions and emissions. With the help of the method, different impact categories can be selected. In the following analysis, the impact of the product life cycle on the "Global Warming Potential 100a" (greenhouse potential) is considered. Using the calculation model, the environmental impacts are expressed in kg CO2.- eq. and assigned to the respective material and energy flows.

Using the calculation model, the environmental impacts are expressed in kg CO2.- eq. and assigned to the respective material and energy flows. The index expresses the warming effect of a specific quantity of a greenhouse gas over a defined period (100 years) compared to that of CO2.Table 3 shows both the percentage distribution over the entire process chain and the quantity data for the individual process steps of both scenarios. In order to obtain a more accurate representation of the respective shares within a process step, the energy or material flows with the largest environmental impacts are additionally listed in the subcategories.The production and associated process steps of 1840 components produced from AlSi10Mg over 8 years using the LPBF process result in total CO2-eq. emissions of 241,396 kg in scenario 1. This results in a value of 131.19 kg CO2-eq per component. At almost 50%, raw material extraction and alloy production entail the largest carbon dioxide emissions.

This results in a value of 131.19 kg CO2-eq per component. At almost 50%, raw material extraction and alloy production entail the largest carbon dioxide emissions. In particular, the energy required to extract primary aluminum predominates with 71,591 kg CO2-eq. The other alloying elements also have a much smaller effect due to their low mass fraction. After the individual raw materials of the alloy are available, electrical energy is again required to produce the alloy. This step causes emissions of 7,404 kg CO2-eq. The subsequent powder production by atomization accounts for 5.14% of the total emissions; here, the electrical energy required clearly predominates. The contribution of the LPBF process is 86,023 kg CO2-eq. Consumables, such as the build platform or silicone lips as well as transport to the post processing (heat treatment etc.) can be neglected. The consumed shielding gas, in this case argon, causes emissions of 9,401 kg CO2-eq.

The consumed shielding gas, in this case argon, causes emissions of 9,401 kg CO2-eq. The value is exceeded by the electrical energy required, which makes a total contribution of 74,299 kg CO2- eq. (30.78%). The subsequent powder preparation and cleaning of the LPBF system have a small contribution of 1.55% and 0.3% respectively. For the heat treatment of the build jobs, too, the electricity demand represents almost the entire contribution of 8,625 kg CO2-eq. Milling, which931separates the component from the build platform, does not have a significant impact on the overall balance (0.69 %).The production of four instead of one component per construction job results in correspondingly higher pollutant emissions (scenario 2). A total of 304,309 kg CO2-eq. is emitted into the environment for the 7360 components. This results in a value of 41.35 kg CO2-eq. per component.

A total of 304,309 kg CO2-eq. is emitted into the environment for the 7360 components. This results in a value of 41.35 kg CO2-eq. per component. At 53.79% (163,676 kg CO2-eq.), raw material extraction and alloy production account for the largest share of the environmental impact in the overall product life cycle. A further breakdown of the process step reveals that the electrical energy used to extract the aluminum in particular predominates at 32.29%. Added to this is a further 8.76% due to the aluminum oxide obtained. The extraction of other alloying elements, as well as the transport of the materials, have a significantly lower influence (< 2.7 %) on the balance. In the powder atomization step, 5.59% of the total amount of CO2-eq. is emitted, which corresponds to 17,026 kg. Also in this process step, the consumed electrical energy dominates with 4.85 %. The influences of the LPBF system and cooling unit remain unchanged in terms of quantity.

Also in this process step, the consumed electrical energy dominates with 4.85 %. The influences of the LPBF system and cooling unit remain unchanged in terms of quantity. The LPBF process causes an emission of 94,874 kg CO2-eq. Also in this case, the electricity demand is the main factor. The resulting environmental impact due to argon consumption amounts to 10,144 kg CO2-eq. Both the powder reconditioning (1.23%), in which the excess powder is sieved, and the cleaning (0.25%) of the LPBF system have only a minor share. The post-processing of the components by means of heat treatment and milling have a share of 2.83 % and 2.18 %, respectively. These process steps are also dominated by the associated power consumption.The following analysis is done for scenario 1, but observations are in accordance with scenario 2. The impact assessment shows that two process steps along the product life cycle dominate the balance for the processing route.

The impact assessment shows that two process steps along the product life cycle dominate the balance for the processing route. On the one hand, the production of raw materials has the greatest negative environmental impact with almost 50 %, corresponding to 119,257 kg CO2-eq. In particular, primary aluminum extraction with its associated electricity consumption has a major impact on the life cycle assessment (29.66%). The main reason for the high electricity consumption is electrolysis. At this point, the first lever for a more sustainable process can already be identified. According to Ostermann [27], if secondary aluminum is used instead of primary aluminum, up to 95% of the energy can be saved in the production phase. Since only aluminum that has already been produced is remelted for secondary aluminum, the electrolysis process is not required. For this, additional monitoring of the material properties must take place at this point to ensure high component reliability.

For this, additional monitoring of the material properties must take place at this point to ensure high component reliability. In addition, the production of the secondary material does not take place exclusively in countries such as China or similar. By using recycled aluminum, the transport routes can be eliminated and resources can be conserved.The second point of the analysis takes a closer look at the LPBF process itself. The emission of 86,023 kg CO2-eq. corresponds to 35.63% of the total emission. As in the case of raw material production, the electrical energy consumed predominates, accounting for 30.78 % of the carbon dioxide equivalent emission. This statement is consistent with the results known from publications (Faludi et al. [21]).932Scenario 1Scenario 2Process stepShare [%] Mass [kg CO2-eq.] Share [%] Mass [kg CO2-eq.]Raw material extraction & Alloying Aluminum production └ 42.84 % └ 29.66 % Electr.

Figure 5 shows the consumption of electrical energy over the entire process chain of the functional unit. 57% of the electricity requirement is incurred for the LPBF process. It should be noted that this percentage is made up of three process steps. The preparation time (flooding the build chamber and heating the build platform to 200 °C) accounts for only 3 % of the total consumption. The manufacturing of the impeller component consumes 13% of the electrical energy and is therefore not the main part of the process chain. The LPBF system and the associated cooling unit consume 42% of the total electricity during the downtime. The downtime is the period between the end of the build job and the start of preparation for the next build job. The process steps whose influence is less than one percent are not shown in the diagram. The total consumption of electrical energy amounts to 238,696 kWh for the defined functional unit.Figure 5.

The process steps whose influence is less than one percent are not shown in the diagram. The total consumption of electrical energy amounts to 238,696 kWh for the defined functional unit.Figure 5. Electrical energy consumption per process step for scenario 1At this point it can already be stated that the downtime of the SLM 280 HL should be kept as low as possible. There are two different approaches to this. Firstly, if the demand is there, an increase in capacity utilization can be considered. Instead of one build job per day, three can be produced per day. The second option is to shut down the plant after each build job and thus remove it from the power grid. This measure could save approx. 97,000 kWh over the 8 years.To illustrate the differences and the associated environmental impacts of the two scenarios, they are compared with each other.

This measure could save approx. 97,000 kWh over the 8 years.To illustrate the differences and the associated environmental impacts of the two scenarios, they are compared with each other. The following Figure 6 illustrates this along the entire process chain.Both raw material extraction and atomization show significant differences between the two scenarios, as more material must be available for four impellers per build job (scenario 2) than for one impeller per build job (scenario 1). Other process steps, such as manufacturing of the LPBF system or powder preparation, show no difference and result in the same amounts of CO2-eq., regardless of the scenario. The CO2 emissions of the LPBF process itself shows significantly smaller deviations between the scenarios than is the case for raw material production. One reason is that a basic consumption of energy is needed in the system regardless of the exposure of the powder.

One reason is that a basic consumption of energy is needed in the system regardless of the exposure of the powder. In addition, there is no difference in cooling, because with the help of the aggregate, the LPBF system is cooled at a certain interval, regardless of the build job. Individual steps in the934preparation process, such as flooding the build area or preheating the build platform, are also independent of the component. The same applies to the recoating process. The only difference between the scenarios is that the exposure time is increased by 2 h 9 min. In the first scenario, the build job is produced with one laser, while in the second scenario, two lasers operate simultaneously to produce four components. This results in a difference of about 8,800 kg CO2-eq. The subsequent post-processing of the components by heat treatment remains the same in terms of emissions.

This results in a difference of about 8,800 kg CO2-eq. The subsequent post-processing of the components by heat treatment remains the same in terms of emissions. In the case of post-processing using a milling machine, the results correspond to the expected differences due to the number of components to be processed.Figure 6: Comparison of CO2-eq. emissions per processing step for scenario 1 and scenario 2In the following step, the two scenarios were optimized with regards to the identified levers. In this case, optimization can be understood as an improvement of existing processes [28]. Secondary aluminum was used instead of primary aluminum and the LPBF system was switched off after each build job. The results of these process chain alterations are shown in Table 4. The first control lever reduced CO2-eq. emissions during aluminum extraction by 94.7%. This corresponds to approximately 99,000 kg CO2-eq. in the first scenario and approximately 135,000 kg CO2-eq.

The first control lever reduced CO2-eq. emissions during aluminum extraction by 94.7%. This corresponds to approximately 99,000 kg CO2-eq. in the first scenario and approximately 135,000 kg CO2-eq. in the second scenario. The second lever showed that regardless of the scenario, about 90,000 kWh, which correspond to about 50,000 kg CO2-eq., could be saved by taking the production system offline. In the overall consideration of the respective scenarios, the levers showed a positive effect in favor of the environmental impact. Total CO2-eq. emissions were reduced from 241,396 kg to 89,947.2 kg in the first scenario. This corresponds to a percentage saving of 63%. The second scenario resulted in a saving of 184,388 kg, or 60.3%.

emissions were reduced from 241,396 kg to 89,947.2 kg in the first scenario. This corresponds to a percentage saving of 63%. The second scenario resulted in a saving of 184,388 kg, or 60.3%. Thus, if the utilization of the build job is maximized, secondary aluminum is used instead of primary aluminum, and the LPBF system is shut down during downtime, this results in an optimized product life cycle for components made of AlSi10Mg, which are manufactured using the LPBF process. This was proven by the amount of CO2-eq. per component. Instead of the 131.19 kg CO2-eq. per component in the first, non-optimized scenario, the emission in the optimized scenario 2 is only 16.29 kg CO2- eq. per component.

energy Powder reconditioning Cleaning Heat treatment Milling Total Mass per part Mass per kg molten powder material Table 4: Overview of total impact assessment for scenario 1 and scenario 2 per process step in163.676,00119.257,0037.149,3620.283,10141.944,00 17.026,07 7.896,25 1.086,56 94.874,81 82.402,806.106,36 22.826,08 7.896,25 1.086,56 46.324,41 33.852,40103.422,00 12.414,72 7.896,25 1.086,56 86.023,52 74.299,504.449,10 12.414,72 7.896,25 1.086,56 33.547,62 21.823,603.730,603.730,543.730,543.730,54768,85 8.625,26 6.623,86 119.920,96 16,29768,85 8.625,26 6.623,86 304.309,00 41,35715,14 8.625,26 1.656,11 241.396,00 131,19715,14 8.625,26 1.656,11 89.947,20 48,88350,42940,45296,39116,80the optimized stateCritical RemarksThe openLCA study, which was carried out within the framework of this work using the LPBF method, multiple simplifications had to be made.

To generate such an extensive process chain with only self-measured data is not possible in this framework for many reasons. In order to be able to represent the product life cycle nevertheless, data from various publications or databases was used.Since there is no usable data on the production of the AlSi10Mg alloy, another data set of an aluminum alloy was modified to the corresponding alloying elements and the assumption was made that the electrical energy required as well as the emissions are the same regardless of the respective alloy. In addition, it is assumed that the raw material is sourced in China and the material is transported to Germany for use. It is not possible to say exactly from which regions of China the individual parts are procured.

It is not possible to say exactly from which regions of China the individual parts are procured. The information on the transport distances is therefore subject to fluctuations and can lead to deviations in the transport energy given.936The powder production process was presented with sources from various publications, which also refer to different materials. Furthermore, only the primary inputs and outputs could be determined. Emissions resulting from chemical reactions during the atomization process could not be accounted for. The powder with a particle size >63 μm was not considered as waste, as it can be used for other processes.Other assumptions that had to be made relate to the equipment used. The LPBF system was accounted for according to Faludi et al. [21]. Although a Renishaw AM250 was analyzed in their work, the assumption was made based on similar weight and size information that the composition is identical.

[21]. Although a Renishaw AM250 was analyzed in their work, the assumption was made based on similar weight and size information that the composition is identical. For the cooling unit and vacuum cleaner, the existing data sets were multiplied by an appropriate factor to arrive at the existing weight of the actual machines used. The electrical energy required to assemble the individual components was not accounted for. This also applies to the disposal of the equipment.Due to the material properties of aluminum according to the LPBF method, it was also necessary to consider heat treatment. However, the electrical energy balanced for this relates to a different component with undefined mass values. The evaporated water during quenching, could neither be determined experimentally nor mathematically, so an assumption had to be made.

The evaporated water during quenching, could neither be determined experimentally nor mathematically, so an assumption had to be made. However, this assumption does not show any relevant effects in the impact estimation.The milling considered in the following is rarely used in reality for finishing and cutting the components. In most cases, the components are reworked using the EDM process or, if necessary, separated from the building platform using a band saw. However, no input and output values could be found in the literature for the EDM process.The previously mentioned assumptions can influence the result and lead to inaccuracies. However, it can be assumed that these inaccuracies are not so large that a completely different result is produced. The analysis has shown that the greatest environmental impact is caused by aluminum extraction (42.8% and 46.6%) and the LPBF process (35.6% and 31.2%).

The analysis has shown that the greatest environmental impact is caused by aluminum extraction (42.8% and 46.6%) and the LPBF process (35.6% and 31.2%). The electrical energy required for the LPBF process was determined using a measuring device (Fluke 435 series II) and therefore represents a reliable value.In general, it can be stated that the LCA results provide guide values that can be used to identify the ecological weaknesses of the LPBF process. The more precisely all product life cycle phases such as manufacturing processes, transport routes, etc. are quantified, the more reliable the LCA result.Conclusion and OutlookBased on this finding, the objective of the present work was first to present the entire product life cycle for LPBF components made of AlSi10Mg from raw material extraction to production. Raw material extraction and alloy production mostly take place in China. Aluminum extraction in particular has proven to be environmentally critical.

Raw material extraction and alloy production mostly take place in China. Aluminum extraction in particular has proven to be environmentally critical. One kilogram of aluminum oxide consumes 13.56 kWh of electrical energy. Silicon extraction requires a similar amount of electricity. Subsequently, the individual raw materials are processed to the desired alloy, which results in an electricity consumption of 1.59 kWh/kg. Since the LPBF process requires the material to be in powder form, powder production by gas atomization was also considered. In this process,937the alloy is melted and pulverized under high pressure. However, the LPBF process requires a certain particle size (20 - 63 μm), which accounts for only 30% of the total yield in atomization. Per kilogram of melt, 8.1 MJ of electrical energy and 0.4 m3 of nitrogen are required. The LPBF system used was balanced using the publication by Faludi et al. [21]. The LPBF process itself is composed of various inputs.

The LPBF system used was balanced using the publication by Faludi et al. [21]. The LPBF process itself is composed of various inputs. In addition to the powder, LPBF requires argon as a shielding gas, a recoater lip and electrical energy. The component can then be removed and the system be cleaned. The powder, which is not melted, is reconditioned and can then be reused. Since in this case the material is an aluminum alloy, it is necessary that the build job is subsequently heat treated. The T6 heat treatment entails a consumption of 25.96 kWh of energy. To complete the manufacturing process, the components are separated from the build platform using milling. This is followed by the individual utilization phase, followed by disposal.The life cycle assessment carried out includes two different utilization models for the build layout.

This is followed by the individual utilization phase, followed by disposal.The life cycle assessment carried out includes two different utilization models for the build layout. In addition to the general product life cycle analysis, the difference between the production of one and four impellers per working day on a building platform over a period of 8 years was examined. The life cycle inventory was prepared using values from experimental measurements, publications and databases. Only the utilization phase was not considered in this work.The 1840 components manufactured in scenario 1 result in total CO2-eq. emissions of 241,396 kg. This results in a value of 131.19 kg CO2-eq. per component. For the 7360 components in scenario 2, a total of 304,309 kg CO2-eq. is emitted into the environment, resulting in a value of 41.35 kg CO2-eq. per component.

per component. For the 7360 components in scenario 2, a total of 304,309 kg CO2-eq. is emitted into the environment, resulting in a value of 41.35 kg CO2-eq. per component. In both scenarios, aluminum extraction dominated with 42.84% and 46.64%, respectively, and electrical energy consumed during the LPBF process dominated with 30.78% and 27.08%, respectively. The influence of individual process chains such as powder production, heat treatment, etc. was in the single-digit percentage range. If the product life cycle phases are considered, it can be seen that, regardless of the different scenarios, the consumption of electrical energy predominates in each phase. Raw material extraction with 28 % and the downtime of the LPBF system with 42 % had the largest influence.

Raw material extraction with 28 % and the downtime of the LPBF system with 42 % had the largest influence. A comparison of the two scenarios with regards to the quantity of emissions emitted shows that the scenario with four components per build job in raw material extraction and atomization has a significantly lower impact on the energy consumption.During the LPBF process, however, this difference was very small. The reasons for this are that many energy flows are independent of the component, such as the basic load of the system, cooling or individual steps in preparation such as flooding of the build area or preheating the build platform. When looking at the electricity consumption, it was also noticeable that 42% and 33% of the electricity demand of the entire product life cycle was attributable to the downtime time of the system for the investigated scenarios.However, the LCA study carried out also has various simplifications.

Since it was not possible to generate a comprehensive product life cycle analysis using only self-measured data within this framework for a variety of reasons, data from different publications or databases had to be used. The assumptions mentioned above can influence the results and lead to inaccuracies. In general, it can be stated in conclusion that the results provide guide values by which it can be seen which ecological weak points the LPBF method has.938One of the results of the present work is that an improved CO2-eq. balance is achieved if secondary aluminum is used for the alloy production. However, no publications could be found that deal with this solution approach for the LPBF process. Accordingly, the use of recycled aluminum raises many questions that still need to be answered before it can be used. One crucial question is whether the material properties of the secondary aluminum are suitable for atomization and subsequent melting by LPBF.

One crucial question is whether the material properties of the secondary aluminum are suitable for atomization and subsequent melting by LPBF. In literature, it is only generally stated that undesirable dross can occur during the preparation of the secondary aluminum. How this affects powder materials for the LPBF process is not known. In addition, recycled materials or reprocessed powder cannot be used in every industry. For powder used in the manufacture of components, for example in the aerospace industry, there are certification processes with high safety requirements regarding powder quality. A first step would be to look at the material properties of secondary aluminum for the LPBF process.

A first step would be to look at the material properties of secondary aluminum for the LPBF process. In addition, a comparison must also be made with conventional manufacturing processes such as casting or milling, since these processes are traditionally used in series production.Furthermore, a consideration of the product life cycle in an overarching cooperation with the manufacturers of the alloy and the powder would be useful in order to generate up-to-date and more specific data, since the technologies are always evolving and the manufacturers themselves have an increasing desire for sustainable production. In addition, companies could help generate a realistic use case for the use phase.With the impact assessment data generated, other impact categories beyond global warming potential can be considered and analyzed. This should be done, as it is possible that process steps that were previously hardly considered can have a strong influence in other impact categories.

This should be done, as it is possible that process steps that were previously hardly considered can have a strong influence in other impact categories. In the evaluation, an already known problem of life cycle assessment becomes apparent. Due to the many different investigation and calculation methods, results from publications are often not comparable with each other. In addition, despite the existing standards, the investigation framework is also calculated with different assumptions and data sets for each analysis. Here it would make sense for a uniform calculation to take place for defined areas.In general, further research focuses on the improvement of component quality and the LPBF method.

 EVALUATION OF THE MECHANICAL BEHAVIOR OF BRONZE..NI COMPOSITES PRODUCED BY SELECTIVE LASER SINTERINGMukesh K. Agarwala, David L. Bourell, Benny Wu, and Joseph J. Beaman. Center for Materials Science and Engineering, and Department of Mechanical Engineering. The University of Texas at Austin, Austin, Texas 78712.ABSTRACT Mechanical properties of Bronze-Nickel composites produced by Selective Laser Sintering (SLS) were evaluated by constant displacement tension tests. These were studied as a function of SLS process parameters - laser power density, scan speed, scan spacing, scan direction and layer thickness. The strength data was then correlated to the microstructure and the part bulk density. To further enhance the part densities and the mechanical properties, post-SLS sintering was studied.

The strength data was then correlated to the microstructure and the part bulk density. To further enhance the part densities and the mechanical properties, post-SLS sintering was studied. The relationships between SLS process parameters, post-SLS sintering parameters and the resulting microstructures, part bulk density and the mechanical properties will be described.INTRODUCTIONMarkets for rapid prototyping continue to grow as the demand for functional parts and prototype tooling increases. Several novel approaches exist to meet this demand.

Several novel approaches exist to meet this demand. In recent years, Selective Laser Sintering (SLS) has emerged as one such process which can, not only produce parts from polymers successfully, but also is capable of producing parts from high temperature materials like metals and ceramics directly, without the aid of any low temperature binders (1,2).Direct SLS of metals is extremely desirable as it avoids the time consuming and often tedious step of binder removal associated with indirect fabrication of metals and ceramics parts. This makes the production of prototype patterns, molds and dies for casting and injection molding, fast and economical.However, any structural part produced, either by direct SLS or any other technique, is of little consequence unless it has the structural integrity to withstand the complex loading conditions that its use may require.

In this study, the mechanical behavior of direct SLS Bronze-Nickel parts were evaluated, as this alloy system is being developed for use as molds in injection molding.EXPERIMENTAL PROCEDURESub-size tension test bars of Bronze-Nickel were made in the high-temperature SLS workstation developed at The University of Texas at Austin (3,4). The dimensions and geometry of the bars used in this study is shown schematically in Figure 1., which conforms closely with the ASTM Standard E8 (5) and the standard MPIF tension test bars for powder metallurgy materials (6).Prealloyed, 90 Wt.% Cu - 10 Wt.% Sn, Bronze powders of particle size 30-50 J,lm and Ni powders of particle size 75-150 J,lm were n1ixed and SLS processed at laser powers of 35 W, 45 W, and 56 W with beam diameters of 0.875 mm, 1.06 mm, and 1.224 mm, respectively. Scan speeds in the range 0.2 to 1.5 inch/second were used. Layer thickness and scan spacing were kept constant at 0.01 and 0.02 inch, respectively.

Scan speeds in the range 0.2 to 1.5 inch/second were used. Layer thickness and scan spacing were kept constant at 0.01 and 0.02 inch, respectively. Scanning was done along the transverse and longitudinal directions of bars and layers were built in the thickness direction.Post-SLS sintering was done for a series of tension test bars processed at 56 W laser power and scan speeds of 0.3"/sec. and 1.5"/sec. Post-SLS sintering of the bars was done at temperatures ranging from 9000C to ll000 C in flowing hydrogen for time periods varying from 1 to 10 hours. All tension testings were done in a constant displacement testing machine at a crosshead speed of 1 mm/min. Densities were measured by direct volume measurement. Phase193- - -::-W----10 -1r-RTFigure 1. A Schematic of Tensile Bar Geometry Used to Make SLS Bronze-Ni Samples. (L=3" , B=O.7", A=l", G=0.75", W=O.2", C=O.5u , R=0.375", and T=0.25")Figure 3.

A Schematic of Tensile Bar Geometry Used to Make SLS Bronze-Ni Samples. (L=3" , B=O.7", A=l", G=0.75", W=O.2", C=O.5u , R=0.375", and T=0.25")Figure 3. SEM of Showing Melting of Bronze With No Melting of Ni.racture urface of SLS Bronze-Nl Part9085~ 80 =- 75 .;; c Q 70--B---- 35W -0 -45W - ~ - 56W-06560o 0.2 0.4 0.6 0.811.2 1.4 1.6Scan Speed (inch/sec)Figure 2. Fractional Density of SLS Bronze-Ni Parts as a Function of Scan Speed and Laser Power.10080---e- 35W -0 -45W~ - 56W.c:.... ~40 s......~r.IJ20oo 0.2 0.4 0.6 0.811.2 1.4 1.6Scan Speed (inch/sec)Figure 4. Tensile Strength of SLS Bronze-Ni Parts as a Function of Scan Speed and Laser Power.analysis and microstructural analysis of the material was done at every step of processing by X-ray diffraction, optical microscopy, EDS and SEM.RESULTS AND DISCUSSIONPreliminary experiments were done to investigate the effect of scau/<lirection and layer thickness.

From these experiments it was found that strengths were considerably lower when scanned in the longitudinal direction when compared to transverse direction. This is due to a better sintered part when the scan vector is short. Nelson, et al (7) have shown that a series of short scan vectors would receive more net energy than a series of longer scan vectors because of less delay between successive pulses.As expected, lowering the layer thickness from 0.02" to 0.01" resulted in higher strength. However, on further lowering the layer thickness to 0.005" resulted in difficulty in spreading fresh layers without disturbing the previously sintered layers. This is especially the case during the early build-up of the part.

This is especially the case during the early build-up of the part. Therefore, all further experiments were done at a constant layer thickness of 0.01" and all scanning was done along the short-transverse direction.OF SCAN SPEED AND LASER POWERAs shown in Figure 2, for a given laser power, density of the SLS Bronze-Nickel parts increased as the scan speed decreases. Also, the density is found to increase with increasing laser power, at a constant scan speed. Higher density is achieved with slower scan speed and higher laser power due to an increased amount of energy input to the powder surface. A higher amount of energy to the powders increases the temperature high enough to result in a larger amount of liquid phase formation. Bronze, which melts incongruently between 8700 C (solidus) and 10300 C (liquidus), exhibits a higher degree of liquid formation as the temperature above the solidus increases.

Bronze, which melts incongruently between 8700 C (solidus) and 10300 C (liquidus), exhibits a higher degree of liquid formation as the temperature above the solidus increases. Also, as temperature increases, the viscosity of molten bronze decreases, facilitating more efficient densification with the solid nickel particles. An increase in temperature also lowers the wetting angle between liquid bronze and solid nickel to result in better wetting characteristics and improved densification. Figure 3 shows the fracture surface of a typical Bronze-Nickel part showing melting of bronze particles while the nickel particles are present without any significant melting. However, at very high laser powers and slow scan speeds (0.2"/sec. - 0.8"/sec.), significant amount of "curling" phenomenon was observed in the parts, as observed by Zong" et.al. also (l).

However, at very high laser powers and slow scan speeds (0.2"/sec. - 0.8"/sec.), significant amount of "curling" phenomenon was observed in the parts, as observed by Zong" et.al. also (l). Experiments with varying bed temperatures showed that "curling" can be minimised by using high bed temperatures while keeping laser power high enough with slow scan speeds necessary for high density parts. However, at bed temperatures above 450°C, the entire powder bed undergoes light sintering and forms a "cake" around the SLS part. This makes it difficult to remove the SLS parts from the surrounding powder bed. Therefore, all experiments in this study were done at a bed temperature of 4500C.Strength, which is primarily a function of fractional density (or porosity) (8), exhibited a similar trend as density with respect to scan speed and laser power, Figure 4. Pores reduce the effective load carrying capacity of a material.

Pores reduce the effective load carrying capacity of a material. Also, pores act as stress concentrators and as effective crack initiation sites. Therefore, a sample with density less than 100% is expected to have a strength less than that of a fully dense wrought material. In addition to fractional density, the strength of such porous compacts depends on various powder characteristics such as particle size, particle size distribution and particle shape. The strength of porous compacts is expected to vary with fractional density p of the compact as follows:(j = C cro f(p)(I)195where (jthe strength of compact, (jo is the wrought strength and f(p) is a fractional densitydependance function. Several studies have attempted to correlate strength with p by various forms of f(p).

Several studies have attempted to correlate strength with p by various forms of f(p). The most commonly used relationsh~p is of the form :(j = C (jo pm(II)where C and m are empirical material constants and p is the fractional density.POST-SLS PROCESSING OF BRONZE-NICKEL PARTSIn an attempt to further densify and hence improve strength, the SLS Bronze-Nickel tensile coupons were further processed by Liquid Phase Sintering (LPS) in hydrogen above the solidus temperature (8700C) of bronze. Before looking at the results of post-SLS processing, it is important to review the basic concepts of LPS (9) to better understand the results, especially with respect to Bronze-Nickel system. Liquid phase sintering in Bronze-Nickel system occurs in three stages:STAGE I: At temperatures above the solidus temperature 8700 C, bronze melts partially and above the liquidus 10300 C it melts completely.

In this stage of LPS, the liquid bronze flows into the pores and a rearrangement of the solid nickel particles takes place. This stage leads to rapid shrinkage and increase in density and hence strength. Most of the densification in LPS is achieved in this stage. If sufficient amount of liquid is formed with low enough viscosity to flow freely, then the final density at the end of stage I is independent of the starting density, which is determined by laser power and scan speed. The amount of liquid phase formed and the viscosity of the liquid formed depends on the sintering temperature. Higher the sintering temperature, higher is the amount of liquid phase fonned with lower the viscosity. But a very high sintering temperature can lead to very high amount of liquid phase with very low viscosity which can result in loss of shape.

But a very high sintering temperature can lead to very high amount of liquid phase with very low viscosity which can result in loss of shape. Particle size of solid phase and rate of heating also determine densification at this stage.STAGE In this stage of LPS, normal densification is accompanied by chemical homogenization between bronze and nickel. Bronze, predominantly copper (90 Wt. %), and nickel homogenize at high temperatures, either by interdiffusion if the homogenization is between solid bronze and solid nickel or by solution-reprecipitation if the homogenization is between liquid bronze and solid nickel, to form a homogeneous solid solution. In either case, the bronze phase disappears completely leaving behind pores in its place and an expanded solid solution of Cu in Ni. This happens due to a faster diffusion rate of Cu into Ni than that of Ni into Cu, which results in the Kirkendall effect. Therefore, the pores created at this stage are referred to as Kirkendall porosity.

Therefore, the pores created at this stage are referred to as Kirkendall porosity. Kuczynski has demonstrated the influence of the Kirkendall effect upon sintering in a series of model experiments (10).' Therefore, in this stage, normal densification is dominated by the growth or swelling of the part due to Kirkendall porosity.:Detailed investigations of LPS have shown that similar material transport mechanisms prevail during stage I as well as stage II. Therefore, a sharp distinction between these stages is not justified. Whether stage I or stage II predominates the process depends on the particular circumstances. Powder particle size, starting density,. sinteringtemperature, time and rate of heating are most influential in determining this.

Powder particle size, starting density,. sinteringtemperature, time and rate of heating are most influential in determining this. Effectof some of these variables on LPS will be discussed laterwithrespecttoSLSBronze-Nickelparts.Figure 5 shows the microstructural changes observed with the different stages of LPS and Figure 6 shows the peak broadening occurring due to formation of a homogeneous solid solution of Cu- Ni.STAGE : Once hon10genization has occurred, only normal shrinkage due to sintering occurs.

But now, the sintering occurs in solid state as the temperature is below the solidus of the newly196of (a) SLS0.3 inch/sec) Showingand Ni as Separate Phases and HighPorosity, (b) Post-SLS Stage II LPS Bronze-Ni (56W, 0.3 inch/sec., 9500C, 1 hour) Showing Nearly Homogeneous Phase of Cu-NiPartSome Residual Bronze and High Porosity,(c) Post-SLS Stage IBronze-Ni Part (56W, inch/sec, 1030oC, 1hour) Showing Nearly of Cu-Ni With SomeLow PorosityI-' ~ co8585(a)(c)8787Cn8989Ni9193Cu-Ni Solid Solution9193959585(b)87899193Figure 6.XRD patterns of (311) peaks of Bronzeand Nickel Showing Peak broadening due toHomogenization (a) SLS Bronze-Ni Part (56W, 0.3 inch/sec) Showing Cu and Ni as Separate Phases (b) Post-SLS Stage II LPS Bronze-Ni I hour) Part Indicating Partial Homogenization of Cu and Ni(56W, 0.3 inch/sec., 950°C,(c) Post-SLS Stage II LPS Bronze-Ni Part (56W, 0.3 inch/sec, 9500C, 6hour) IndicatingComplete Homogenization of Cu and Ni.95formed solid solution of Cu-Ni (-11000 C).

Therefore,even for prolonged periods of times, shrinkage or densification during this stage is very small, leaving an overall expansion or shrinkage in the samples, depending on whether first stage or second stage of LPS dominates.EFFECT OF SINTERING TEMPERATURE AND TIMEAs shown in Figures 7, at temperatures below 10000 C the density and strength of SLS parts decrease irrespective of starting densities. The density and strength reach a minimum and then increase slowly to nearly starting densities. At such low temPeratures, amount of liquid phase formed is very small with high viscosity which prevents any significant densification normally associated with stage I LPS. Instead, stage II dominates the LPS initially and then stage III continues till the end of the process.At 1060°· C, a temperature well above the liquidus of bronze, bronze is completely in liquid state and has sufficiently low viscosity to allow stage I LPS to dominate.

Therefore, the density and hence the strength of the SLS parts increase rapidly at short times of 1 hour. Following this rapid gain in density and strength, there is a slight loss of density and strength due to stage TILPS when the density and strength decrease due to Kirkendall porosity. Some of this loss in density and strength is regained in stage TIl of LPS due to solid state sintering of Cu-Ni solid solution. At such a high temperature the flow of liquid in stage I is free of any restriction due to low enough viscosity. Therefore, the final density and strength at the end of stage I is independent of starting density and strength. At temperature above 10600 C, the SLS parts exhibited loss of shape even at small times due to significantly lower viscosity of the liquid bronze.At the liquidus temperature of 1030° C, liquid bronze is more viscous than at 10600 C. Therefore, flow of liquid into pores and rearrangement of solid nickel particles is less dominant.

Therefore, flow of liquid into pores and rearrangement of solid nickel particles is less dominant. Therefore, the density and strength in stage I at 1030° C is lower than that at 10600 C. In a high starting-density part, the rigid network of solid nickel particles restrict the flow further and prevent their rearrangement. While in a lower-starting-density part the liquid is more free to flow and rearrange the solid particles. Also, the interdiffusion lengths in a high-starting-density part is smaller, thereby promoting stage II LPS and thus resulting in lower final strengths and density than that for a lower-starting-density part. This effect is more pronounced at 1000° C when the stage II homogenization completely dominates over the stage I densification for a higher-starting- density part.

This effect is more pronounced at 1000° C when the stage II homogenization completely dominates over the stage I densification for a higher-starting- density part. Therefore, at 10000 C, while a high-starting density part exhibits worsening of density and strength, a low-starting-density part exhibits improvements in density and strength.ANISOTROPIC LINEAR DIMENSIONAL CHANGESNet volume changes (expansion or shrinkage) of SLS parts during LPS was in accordance with the observed density and strength changes. However, the linear dimensional changes, &/Lo, was not same in all three "directions (11). They exhibited some rather interesting trends reflecting the anistropy in SLS parts.In general, parts that exhibited swelling or growth during LPS, had a net volume expansion exceeding 10% and was mostly of the order of 15%, resulting in a decrease in fractional density of In accordance with stage II LPS, this decrease was the order of 8% to 12%, as seen earlier.

observed over a relatively short period of time followed by a nearly constant density, as would be expected in stage III of LPS. However, the linear dimensional expansion in these parts was anisotropic with least expansion (-2%-3%) in the thickness direction, Le. direction normal to the sintering plane. Most of the expansion (5%-9%) was confined to the transverse direction (scan direction), Figure 8 (a). A higher degree of sintering along the scan direction, during SLS, results in better density along the transverse direction, thus forming a rigid network of solid nickel and decreasing the interdiffusion lengths. Therefore, stage I LPS is suppressed and stage II LPS dominates along the scan direction yielding a high linear expansion. While a lower degree of199350300250 CO ~6200.c:... CI) 150 I: I1J... C;; 100--_EJ50o 01234 5 Time (Hours)61009590-~ 85-~80....fI).

-1000oe - - -.to - - 9500e<9-sintering normal to the sintering plane results in poor density between layers leaving a loose network of solid nickel and large interdiffusion lengths. Therefore, stage I LPS is less suppressed and stage II LPS is less dominating between the layers resulting in considerably lower linear expansion. Also, flow of liquid bronze in stage I along the thickness direction is aided by gravity as the samples were post processed with the thickness direction along the direction of gravity.the parts exhibiting overall shrinkage or densification also exhibited volume changes according to changes in density, but the linear dimensional changes in the three directions were different. As explained above, since stage I LPS is more dominant along the thickness direction than along the scan direction, higher linear shrinkage was observed in the thickness direction than along the transverse direction.

Shrinkage in the longitudinal direction was comparable but higher than along the transverse direction, Figure 8(b).Like the parts which exhibited growth,STRENGTH VARIAnON WITH DENSITYAs discussed earlier, pores reduce the effective load carrying capacity of a material, thus decreasing the strength of the material. Equation II relates the strength to the fractional density empirically, based on the premise that pores reduce the load bearing capacity of the material. Based on the discussions of results thus far, it can be seen that both the strength as well as density of SLS parts increased with increasing laser power and decreasing scan speed. Therefore, it seems reasonable to relate the strength and density of SLS parts irrespective of the laser power or scan speeds at which they were processed.

Therefore, it seems reasonable to relate the strength and density of SLS parts irrespective of the laser power or scan speeds at which they were processed. As shown in Figure 9, a near perfect fit was obtained for the strength and density by relating them using the power law Equation II, with the empirical constants Ccro and m as 590.8 and 9.43 respectively. A high exponent factor, m, of 9.43 indicates large gains in strength for small incremental gains in density. This makes it possible to produce less than fully dense yet structurally sound parts of Bronze-Nickel by direct SLS, without any post-processing. Small incremental gains in density are achievable by increasing the laser power andlor decreasing the scan speed.

Small incremental gains in density are achievable by increasing the laser power andlor decreasing the scan speed. The "curl" phenomenon associated with high laser power and slow scan speeds can be overcome by using higher bias temperatures in the high-temperature workstation, if light sintering of the powder bed can be avoided or minimised.Liquid Phase Sintering (LPS) of the SLS Bronze-Ni parts results in an increase or decrease in density and strengths simultaneously as a function of LPS time and temperature, as discussed earlier. Therefore, relating strength with density using the power law form given by Equation II, yields a near perfect fit with the empirical constants Ccro and m as 527.94 and 8.73 respectively, Figure 9(b). From comparison of the empirical constants Ccro and m for SLS Bronze-Nickel parts and for post-SLS LPS Bronze-Nickel parts, it is clear that density and strength vary in a similar way irrespective of the processing.

During SLS scan speed and laser power determine the density and strength, while during LPS time and temperature determine the density and strength. SLS laser power is analogous to the LPS temperature as it determines the processing temperature. Similarly, SLS scan speed is analogous to the LPS time as it determines the time of processing. Therefore, by careful manipulation of these four variables it is possible to produce high density, structurally sound parts with no "curl" by direct SLS followed by LPS in very short processing time periods.CONCLUSIONSDirect SLS of Bronze-Nickel parts was studied by evaluating the density and strength of the parts as a function of SLS process parameters: laser power, scan speed, scan direction, and layer thickness. Post-SLS processing of the parts was done by Liquid Phase Sintering (LPS) to further enhance the density and strength. LPS of the SLS parts was studied as a function of time and temperature.

4. LPS of SLS Bronze-Nickel parts improved density and strength at temperatures well above the liquidus temperature of bronze. 5. By LPS at temperatures well below the liquidus temperature of Bronze, density and strength of SLS Bronze-Nickel parts decreased due to Kirkendall porosity resulting from the homogenization of Bronze and Ni. 6. Linear dimensional changes during LPS were anisotropic. 7. Strength anq density of post-SLS LPS Bronze-Nickel parts show a similar power law dependence as the SLS Bronze-Nickel parts.ACKNOWLEDGMENTS The authors gratefully acknowledge research grants from.DTM Corporation, Austin, Texas and NSF Small Grant for Exploratory Research, Grant # DDM-9312603.REFERENCES 1. G.Zong, Y.Wu, N.Tran. I.Lee, D.L.Bourell, J.J.Beaman, and H.L.Marcus, "Direct Selective Laser Sintering of High Temperature Materials," Proc. of the SFF Symp., Austin, Texas, Aug 5, 1992. 2.

G.Zong, Y.Wu, N.Tran. I.Lee, D.L.Bourell, J.J.Beaman, and H.L.Marcus, "Direct Selective Laser Sintering of High Temperature Materials," Proc. of the SFF Symp., Austin, Texas, Aug 5, 1992. 2. D.L.Bourell, H.L.Marcus, J.W.Barlow, and J.J.Beaman, "Selective Laser Sintering of Metals and Ceramics," Int. J. Powder Met., 28(4), 369, 1992. 3. S.Das, J.C.McWilliams, Y.Wu, and J.J.Beaman, "Design of a High Temperature Workstation for the Selective Laser Sintering Process," Proc. of the SFF Symp., Austin, Texas, Aug, 1991. 4. J.McWilliams, C.Hysinger, and J.J.Beaman, "Design of a High Temperature Process Chamber for the Selective Laser Sintering Process," Proc. of the SFF Symp., Austin, Texas, Aug 3-5, 1992. 5. ASTM, "Standard Methods of Tension Testing of Metallic Material," 1977 Annual Book of ASTM Standards, Part 6 : Copper and Copper Alloys, ASTM Designation: E8-77, 898~917, 1977.. 6. Metals Handbook, Vol.7, 9th Ed., American Society of Metals, Metals Park, OH. 7.

6. Metals Handbook, Vol.7, 9th Ed., American Society of Metals, Metals Park, OH. 7. J.C.Nelson, and J.W.Barlow, "Relating Operating Parameters Between Selective Sintering Machines which have Different Scanner Geometries and Laser Spot Sizes," SFF Symp., Austin, Texas, Aug 3-5, 1992. 8. R.M.German, "Powder Metallurgy Science," Metal Powder Industries Federation, Princeton, NJ. 9. F.V.Lenel, "Powder Metallurgy - Principles and Applications," Metal Powder Industries Federation, Princeton, NJ. 10. G.C.Kuczynski, "Model Experiments and the Theory of Sintering." From "Sintering Papers," Ed. by S.Somiya and Y.Moriyoshi, Elsevier Applied Science, London and New York. 11. J.C.Nelson, N.K.Vail, M.M.Sun, and J.W.Barlow, "Post-Processing of Selective Laser Sintered Polycarbonate Parts," Proc. of the SFF Symp., Austin, Texas, Aug, 1991.203of the.

ALUATIONS OF TEMPERATURE MEASUREMENTS BY NEAR-INFRARED THERMOGRAPHY IN POWDER-BASED ELECTRON-BEAM ADDITIVE MANUFACTURINGSteven Price*, Ken Cooper**, Kevin Chou*Department of Mechanical Engineering The University of Alabama, Tuscaloosa, AL 35487, USA**Advanced Manufacturing Technology Team Marshall Space Flight Center, Huntsville, AL 35812, USAAbstractPowder-based electron beam additive manufacturing (EBAM) has gained increased usage in different industries. Process monitoring such as temperatures may offer important information. However, temperature measurements in EBAM are challenging because of high temperature ranges, extreme gradients and fast transient response.In this study, temperature measurements during the EBAM process, in particular, around the electron beam scanning area, were attempted using a near-infrared thermal camera. The obtained temperature data demonstrated the feasibility of such a measuring technique.

The obtained temperature data demonstrated the feasibility of such a measuring technique. The thermal camera was able to capture the pre-heating, contour melting, and hatch melting events. Further, the achievable spatial resolution is around 12 µm when using a 350 mm lens.IntroductionAdditive manufacturing (AM) based on “layer-by-layer” fabrications is a fast-growing emerging technology, by which physical solid parts are made directly from electronic data, generally files from computer-aided design (CAD) software. This group of technologies offers many design and manufacturing advantages such as short lead time, part geometry complexity, and no tooling, etc. Electron Beam Additive Manufacturing (EBAM) is one of relatively new AM technologies. Similar to electron-beam welding, EBAM utilizes a high-energy electron beam to melt and fuse metal powders and build parts in an additive fashion.

Similar to electron-beam welding, EBAM utilizes a high-energy electron beam to melt and fuse metal powders and build parts in an additive fashion. Moreover, EBAM is one of a few AM technologies capable of making full-density metallic parts and has drastically extended AM applications for direct digital manufacturing of structural components. There are two kinds of EBAM technologies, different in raw material stock: powder-bed vs. wire-feed. The latter one, developed by NASA’s Langley Research Center, is known as electron beam freeform fabrications (EBF3) in industry, and is mainly utilized for repair and adding common features to a large structure. On the other hand, the powder-based EBAM technology, commercialized by Arcam AB, has been explored for more net-shape applications in various fields such as aerospace, biomedical, and automotive, etc.

The subject of this paper is the powder-based EBAM and will be referred simply as EBAM through the paper for the sake of space saving.761Despite of EBAM’s capability and potential benefits, applications using such a technology are still not widespread. Part quality from EBAM is sensitive to the process parameters, and yet the setting from the manufacturer may not be necessarily optimal for different part geometries. Further, part quality variations have been a major factor that hinders the certification of EBAM parts for critical applications. Therefore, the effect of process parameters on part attributes need to be studied in order to efficiently implement EBAM applications. However, experimental testing is time consuming and costly. On the other hand, process physics modeling may offer a more efficient means to tackle such an issue.

However, experimental testing is time consuming and costly. On the other hand, process physics modeling may offer a more efficient means to tackle such an issue. EBAM is a rather complicated process and fundamental understanding of process physics is also a key to improve the process performance and part quality. However, there is rather limited modeling work for EBAM reported in literature. Despite of increasing studies in the EBAM process, most have been for applications, case studies, or microstructures/properties, etc. Zäh and Lutzman developed a simplified heat transfer model of EBAM to study the scan speed and beam power effects on the weld-pool geometry [1]. Shen and Chou recently initiated a study to investigate the thermal phenomenon of the EBAM process.

Shen and Chou recently initiated a study to investigate the thermal phenomenon of the EBAM process. A finite element (FE) model incorporating Gaussian heat flux distribution, fusion latent heat, temperature-dependent thermal properties was developed to study the thermal response in the build volume when subject to a moving heat source of high intensity [2].On the other hand, EBAM process measurements are important as well for model validations and process monitoring, but, fairly rare in literature. Zäh and Lutzman applied thermocouples, attached to the build plate, to evaluate temperature response during the EBAM process to evaluate the model developed by this research group [1]. However, because of very high temperatures and steep gradients, thermocouple techniques are limited in accuracy and resolutions. As EBAM, most AM technologies are thermally dominant processes and temperature distributions and history are critical to the process characteristics and performance.

As EBAM, most AM technologies are thermally dominant processes and temperature distributions and history are critical to the process characteristics and performance. There have been different approaches attempted for temperature measurements in AM processes. Most studies used an infrared thermal camera to acquire temperature data, while a few studies used thermocouples to measure point temperatures. Some of the infrared cameras used included a Raytheon radiance HS infrared camera [3, 4], a Thermaviz two-wavelength pyrometer [5], and a FLIR Systems Prism DS digital infrared camera [6]. A Mikron M67S infrared temperature sensor [7] and a Raytech MR1S dual wavelength instrument [8] were also used to measure the average temperature at the center of the molten pool. Most of the cameras had a spectral band pass filter between 3 and 5 µm, although a longer range of 8 to 12 µm [9] and a short wavelength of 0.7 to 1.06 µm [10, 11] were also used.

Most of the cameras had a spectral band pass filter between 3 and 5 µm, although a longer range of 8 to 12 µm [9] and a short wavelength of 0.7 to 1.06 µm [10, 11] were also used. The frame rates used ranged from 1 frame every 2 seconds [5] to 300 frames per second [10]. The cameras were either mounted exterior to the process equipment or had a fiber optic cable mounted coaxially on a head inside a machine. A majority of the thermal cameras were calibrated using an ultrahigh temperature blackbody calibrator and a thermocouple. One camera was calibrated using a NIST tungsten filament [5]. An experimental apparatus used a high shutter-speed camera and a 337 nm near ultraviolet (UV) laser in conjunction with the infrared camera [10, 11]. Griffith et al. [6] used a 12-bit digital camera with a telephoto lens and a 650 nm broad band filter to capture images of a melt pool.

Griffith et al. [6] used a 12-bit digital camera with a telephoto lens and a 650 nm broad band filter to capture images of a melt pool. According to the temperatures recorded with this method, the results were similar to thermocouple and infrared measurements except for those inside the melt region of the melt pool.762For thermocouple applications, Yang et al. [12] sandwiched a K-type thermocouple between metal tapes to record temperatures in an ultrasonic consolidation process. Sun et al. [13] and Pasandideh-Fard et al. [14] both embedded K-type thermocouples into the material deposition base of studied AM processes. Sun et al. studied the thermal cycles a part goes through during a fused deposition modeling process. On the other hand, Pasandideh-Fard et al. studied the solidification of tin droplets on a stainless steel plate. Both Yang et al. and Sun et al. used a sampling rate of 1000 Hz [12,13]. Besides, Griffith et al.

studied the solidification of tin droplets on a stainless steel plate. Both Yang et al. and Sun et al. used a sampling rate of 1000 Hz [12,13]. Besides, Griffith et al. [6] inserted C-type thermocouples into a build part during a laser engineered net shaping (LENS) process in order to study the severity of the thermal cycles caused by the process.Measuring temperature has always been an interesting subject in most manufacturing process studies as it provides valuable insight of process details. However, temperature measurements can be difficult for different reasons, e.g., temperature range, material issues, and transient response, etc. It is especially challenging in some high powder, localized and high rate processes such as EBAM. For high temperature processes, non-contact measurements such as infrared are preferred.

It is especially challenging in some high powder, localized and high rate processes such as EBAM. For high temperature processes, non-contact measurements such as infrared are preferred. However, because of x-ray generations in EBAM processing, a leaded glass is attached on the viewport of the machine, and the glass may affect the transmission, In general, the transmission rate of a leaded glass for the infrared bandwidth can be very low, < 60% [15]. On the other hand, for the wavelength from 500 nm to 2 µm, the transmission rate of most glasses (transparent objects) is less sensitive. Hence, near-infrared (NIR) thermal cameras may be more suitable for such a task. Indeed, NIR thermal cameras have been employed for temperature measurement studies in different manufacturing processes.Haferkamp et al. [16] used a CCD camera with a spectral range up to 1.3 µm to measure the temperature of cast aluminum alloy parts during solidification.

[16] used a CCD camera with a spectral range up to 1.3 µm to measure the temperature of cast aluminum alloy parts during solidification. A mold made of IR- transparent silica aerogel was used in order to capture NIR images of the cast part while it was in the mold. Ageorges et al. [17] used a Photometrics Sensys CCD Camera to measure the temperatures of the heating element during a resistance welding process. The camera detected wavelengths up to 0.9 µm. Temperatures from 180 °C to around 500 °C were in the measuring range of the camera. Outeiro et al. [18] also used a CCD camera to study the temperatures of the tool in a 3D cutting process. The camera used an 850 nm high pass filter. The temperature range of the camera was dependent on the length of the exposition time of the camera. The camera’s minimum temperature ranged from 420 to 542 °C and its maximum temperature ranged from 549 to 757 °C.M’Saoubi et al.

The camera’s minimum temperature ranged from 420 to 542 °C and its maximum temperature ranged from 549 to 757 °C.M’Saoubi et al. [19] used a CCD senor based NIR imager with a spectral range of 0.85 to 1.1 µm to measure the tool temperature during a turning process. They compared the tool temperatures for various coatings, feed rates, and cutting speeds. In a continued study, the same research group [20] used the same NIR imager to measure the tool temperature during an orthogonal cutting process. The workpieces were made of carbon steel AISI 3115 and stainless steel AISI 316L. The effects that various heat treatments on the workpiece have on the tool temperatures were analyzed using the NIR images. Rech [21] conducted a similar experiment where the author used a CCD infrared camera to measure the tool temperature during a turning operation. The workpiece material was AISI 4142H steel. The experiment examined the tool coating effects on the tool temperatures.

The workpiece material was AISI 4142H steel. The experiment examined the tool coating effects on the tool temperatures. The camera had a detectable range of 500 to 1000 °C. The maximum temperature measured on the tool was 653 °C. Jackson and Gossard [22] used a763Canon 350D to measure the substrate temperature in a molecular beam epitaxy system. The spectral range of the camera was 720 to 1100 nm and the camera was able to measure the substrate temperatures in the range of 500 °C to 515 °C. Craeghs et al. [23] examined the effects of various factors during a laser melting process using a high speed NIR camera and a photodiode sensor to determine the temperatures of the melt pool. A 700 to 950 nm low pass filter was used for the camera. Mirrors were used to reflect the image of the process to the camera.

A 700 to 950 nm low pass filter was used for the camera. Mirrors were used to reflect the image of the process to the camera. The NIR images were analyzed to detect the edge of the melt pool by determining which isotherm corresponded with the edge.In this research, an NIR thermal camera was tested to attempt temperature measurements in the EBAM process. The objective of the current phase was to evaluate the feasibility and capability of an NIR thermal camera for investigating the thermal characteristics of the EBAM process. The overall goal of this research is to achieve temperature measurements in the EBAM for process model validations and process monitoring, also to investigate process parameter effects on the thermal characteristics of the EBAM process.Experimental SetupThe EBAM machine used for temperature measurements was an S12 system, from Arcam, at NASA’s Marshall Space Flight Center (MSFC).

Further, it was decided to acquire data images through the 127 mm diameter viewport at the front side of the machine, Figure 1. In addition to the leaded glass mentioned above, there is also a high-strength vacuum glass attached to the viewport, which will affect the radiation transmission as well at certain wavelengths. Thus, an NIR thermal camera, MCS640 from LumaSense, was used. The camera has a spectral range of 780 nm to 1080 nm wavelength, has a 640 by 480 uncooled focal plane array (FPA) sensor (amorphous Si based). The detectable temperature range is from 600 °C up to 3000 °C, with different ranges of calibrations. The maximum frame rate of the image capturing system is at 60 Hz. The general specification of the camera is listed in Table 1 below.Table 1.

The maximum frame rate of the image capturing system is at 60 Hz. The general specification of the camera is listed in Table 1 below.Table 1. Specification of the NIR thermal camera used.Wavelength (nm)FPAMax frame rate (Hz)Temperature range (°C)780 – 1080640 by 48060600 - 3000Two types of lenses, with different focal distances, were employed for different purposes. One lens (A) has a minimal 500 mm working distance and a view area of 31 mm by 23 mm at 500 mm distance (FOV angle: 3.5° by 2.6°). The other one (B), a closed-up lens, has a nominal 350 mm working distance and with a view area of only 5.3 mm by 4.0 mm (FOV angle: 0.88° by 0.65°). Lens A results in a spatial resolution about 48 µm in both horizontal and vertical directions at 500 mm distance. On the other hand, Lens B gives a finer spatial resolution, approximately 8.3 µm.

Lens A results in a spatial resolution about 48 µm in both horizontal and vertical directions at 500 mm distance. On the other hand, Lens B gives a finer spatial resolution, approximately 8.3 µm. The thermal camera was calibrated in 3 different temperature ranges, for Lens A: (650 - 900, 850 - 1250, 1100 - 1800 °C), and for Lens B: (750 - 1100, 915 - 1400, 1320 - 2300 °C) using a black body. For Lens B, the calibration was conducted with and without glasses (identical to those used in S12) to estimate the overall transmission rate of the applied glasses. The transmission rate has been determined to be about 0.54. Emissivity was not known when the testing was conducted. Therefore, a value of 0.3 (for metal) was assumed; however,764this could probably be deviated from the actual values because of the powder material used and the high temperature condition involved.

Further, the integration time for imaging is 16.25 ms, 1.7 ms, 50 µs for the temperature range 1 (low), 2 (medium) and 3 (high), respectively. The data acquisition system included a network cable and a laptop, connected via a gigabit Ethernet express card.The S12 machine, thermal camera setup using a sturdy tripod, and the heat shield covering the build area are shown in Figure 1. Note that the heat shield that covers the build area has the slit area, where the view is made possible. The camera lens was set close to the glass window, straight targeting the build plate area. Figure 2 is a sketch showing the position and orientation of the camera, relative to the machine. The estimated vertical tilt angle was to be about 55°, no horizontal yaw or roll angle.Figure 1.

Figure 2 is a sketch showing the position and orientation of the camera, relative to the machine. The estimated vertical tilt angle was to be about 55°, no horizontal yaw or roll angle.Figure 1. Left: EBAM machine, Center: thermal camera setup, and Right: heat shield.EBAM MachineEBAM MachineNear Infrared CameraHeat ShieldViewportBuild AreaHeat ShieldViewportNear Infrared CameraBuild AreaSide ViewTop ViewFigure 2. A sketch showing position and orientation of the camera relative to the EBAM machine (drawing not to scale).The CAD model designed and built in the machine for temperature measurements using the Lens A was just a simple 25.4 mm square block. While the CAD model built for measurements using the Lens B was several thin-walled structures, shown in Figure 3. The length and width of each thin wall is 25.4 mm and 3.8 mm, respectively. The overall height of the model was 25.4 mm as well.

The length and width of each thin wall is 25.4 mm and 3.8 mm, respectively. The overall height of the model was 25.4 mm as well. The model was designed to evaluate the spatial resolution against the specification. The EBAM process parameters used including the beam voltage, offset765current, beam speed, etc., were default values typically employed for Ti-6Al-4V powders under normal geometry and operations.Figure 3. Build part in EBAM when measuring temperatures using Lens B.Results and DiscussionThe heating cycles in the EBAM process have several phases including build platform (start-plate) heating, preheating, contour melting and hatch melting, etc., and in each phase, some process parameters may be very different. First, once the required vacuum level is achieved, heating of the build platform (~10 mm thick stainless steel plate) is initiated, using a single beam scanning at a very high speed.

First, once the required vacuum level is achieved, heating of the build platform (~10 mm thick stainless steel plate) is initiated, using a single beam scanning at a very high speed. The default target temperature level is around 730 °C (measured by a thermocouple attached to the bottom of the plate). During this phase, the NIR thermal camera was not able to detect temperature rise until a later stage. It was also noted that the heating is paused at about 400 to 450 °C (plate bottom temperature) for about 10 to 20 minutes for heat shield out-gassing.A few sections below describe the build process and thermal images acquired when using Lens A, in making the 25.4 mm square block in the EBAM machine. The build cycle starts when the start-plate bottom temperature reaches 730 °C. During each cycle (each layer fabrication), the process involves powder spreading (via a raking mechanism), powder preheating, contour melting, and hatch melting.

During each cycle (each layer fabrication), the process involves powder spreading (via a raking mechanism), powder preheating, contour melting, and hatch melting. During the powder spread, a metal rake is applied to distribute one layer of powders, 100 µm in this test. The build part surface temperature is instantaneously reduced, to a level below what the thermal camera can sense, since the powders spread from the hopper have a much lower temperature.In the next phase, powder preheating cycle starts using also a single beam at a high speed (visually showing multiple scanning line pattern), and involves multiple scanning passes (alternative raster pattern at different layers) to reach a uniform high temperature across the entire powder bed surface. Row 1 of Figure 4 shows examples of NIR thermal images during preheating at Low and Medium Temperature Ranges, respectively. The line pattern of high temperature zones is noticed, as high as over 1000 °C.

The line pattern of high temperature zones is noticed, as high as over 1000 °C. Another observation to point out is that, recalling the camera has to see through the slit area of the heat shield (Figure 1, Right photo), the seemingly vertical bands observed from the Low Temperature Range represent the areas blocked by the uncut portion of the heat shield.766Following the preheating, the contour melting takes place, during which multiple split beams move across the powder-layer surface tracing the boundary defined by the cross-section at that specific layer (a square in this case) of the build model. It was observed that there appears to have 2 stages in this phase. During the first stage, several split electron beams quickly pulse statically around the part boundary of that layer, starting to melt and fuse the powders together. Figure 4, Row 2 images show examples of NIR images occurred in the first part of contour melting collected at 3 Temperature Ranges.

Figure 4, Row 2 images show examples of NIR images occurred in the first part of contour melting collected at 3 Temperature Ranges. The boundary of the model at that specific layer can be clearly observed, evidenced by the high temperature spots. The distance (in pixel number) between the left and right borders was used to calculate the corresponding spatial resolution, estimated to be 47.6 µm, consistent to the spec. Note that due to the vertical tilt angle of the camera setting, the dimension along the vertical direction is slightly distorted. During the second stage of the contour melting, several electron beams (split) outline the outer boundary more slowly, causing completely melting of the contour.

During the second stage of the contour melting, several electron beams (split) outline the outer boundary more slowly, causing completely melting of the contour. Figure 4, Row 3 images show examples of NIR images occurred in the second stage of contour melting, also at 3 Temperature Ranges.The final phase of each layer build cycle is the hatch melting, during which a single beam scans through the inside of the contour, with a straight rater pattern. The scanning speed is lower than those used during the preheating and contour melting, though visually still quite fast. Figure 5 shows examples of NIR images acquired during the hatch melting, again with 3 Temperature Ranges. It is recognized that the thermal behavior in the hatch melting resembles the features of common moving heat source applications, a localized high temperature zone followed by a tail with rapid temperature decay, opposite to the moving direction.

The temperature range is very high, with the peak well above the melting point of Ti-6Al-4V, 1650 °C.The thermal images observations during the build using Lens B, same setup as using Lens A except a shorter distance to the target, were quite similar to those in the square-block build with Lens A. They followed the various heating phases described earlier, however, with a smaller view area. Since the purpose of this particular testing was to evaluate the spatial resolution of the thermal camera and lens, this section just describes the contour melting stage, during which the spatial resolution was determined. Figure 6 shows the NIR image example during contour melting (first stage) in the area including one of the thin wall columns, horizontal orientation. The vertical distance between the high-temperature zones (centers) was used to estimate the spatial resolution, estimated to be about 16 µm. Due to the vertical tilt angle, note that the resolution is distorted.

Due to the vertical tilt angle, note that the resolution is distorted. On the other hand, the images taken during contour melting of the vertical thin-wall sections gave a spatial resolution of 12 µm. Another observation is that because of a much smaller area in the scanned cross-section, the hatch melting does not clearly show the moving heat source phenomenon illustrated in previous sections. The fast scanning combined with the short transverse distance to travel results in such a result.The acquired thermal images/videos offer visualization to interpret the EBAM process. In addition, the data collected can be analyzed to obtain more insight of the process phenomena. Two examples are discussed below. Figure 7 below shows an example of the melt-pool shape and size during the hatch melting (square block building).

An example of the melt pool shape and size.Figure 8 below shows an example of a temperature profile across the melt pool center, along the electron beam moving direction during the hatch melting, at High Temperature Range. Data from multiple frames, i.e., electron beams at different locations and different files was analyzed. Both of the average (blue line) and standard deviation (vertical bars) values are shown. The result shows good repeatability of the temperature distributions along the beam moving direction. Again, the profile exhibits the thermal behavior similar to a moving source of heat phenomenon: rapid temperature rise at the heat source front, and a quick decay in the back. Further, the phase change region can also be clearly identified as the plateau at around 1450 C, which should correspond to the melting point of Ti-6Al-4V.

Further, the phase change region can also be clearly identified as the plateau at around 1450 C, which should correspond to the melting point of Ti-6Al-4V. Note that, as explained earlier, the emissivity of Ti powders and the NIR transmission ratio of the glasses were both assumed (Lens A), mostly inaccurate, during this experiment. The temperature profile may offer some hints for the emissivity estimate.)C ° (e r u t a r e p m e T2000 1900 1800 1700 1600 1500 1400 1300 1200 1100 100001000 2000 3000 4000 5000 6000 7000 8000Distance (µm)Figure 8. Temperature profile across the melt pool, along the electron beam moving direction (Average data with standard deviation bars).770ConclusionsTemperature measurements during the EBAM process have been attempted using an NIR thermal camera incorporating 2 different lenses individually. Two different CAD models were fabricated separately for acquiring thermal images during the build in the EBAM machine.

Two different CAD models were fabricated separately for acquiring thermal images during the build in the EBAM machine. The feasibility and capability have been evaluated. The major results are summarized as below.Temperature measurements in the EBAM process by an NIR thermal camera are feasible; the camera is able to capture different events at different temperature ranges.Spatial resolutions, lens dependent, were confirmed by images collected during the build of particularly designed models, about 12 µm when using a closed-up lens.The temperature measurements in the hatch melting are repeatable, as evidenced in Figure 8.  One of the challenges is to determine the emissivity of Ti powders at high temperatures.  Future work will include comparison with the simulations for model validations and studying process parameter effects on the thermal characteristics in the EBAM process.AcknowledgementThis research is supported by NASA, No.

NNX11AM11A, and is a collaboration between The University of Alabama and MSFC. Support from Majid Babai, Chief of Nonmetallic Branch at MSFC, is greatly appreciated.References[1] Zäh M. F. and Lutzmann S., (2010), “Modelling and simulation of electron beam melting,”Production Engineering, 4(1), pp. 15-23.[2] Shen, N. and K. Chou, (2012), “Thermal Modeling of Electron Beam Additive Manufacturing Process – Powder Sintering Effect,” ASME International Manufacturing Science and Engineering Conference, Notre Dame, IN, June 4-8, 2012 (MSEC2012-7253).[3] Fischer, P., Locher, M., Romano, V., Weber, H. P., Kolossov, S., Glardon, R., 2004, “Temperature measurements during selective laser sintering of titanium powder,” International Journal of Machine Tools and Manufacture, 44, pp.

M., 2001, “Geometry Modeling and Control by Infrared and in Thermal Manufacturing with Material Deposition,” Journal ofLaser Sensing Manufacturing Science Engineering, 123, pp. 45-53.[10] Hu, D., Kovaceic, R., 2003, “Modeling and measuring the thermal behavior of the molten pool in closed loop controlled laser-based additive manufacturing,” Proceedings of the Institution of Mechanical Engineers, Part B: Journal of Engineering Manufacture, 217, pp. 441-452.[11] Hu, D., Kovaceic, R., 2003, “Sensing, modeling and control for laser based additive manufacturing,” International Journal of Machine Tools and Manufacture, 43, pp. 51-60.[12] Yang, Y., Janaki Ram, G. D., Stucker, B. E., 2009, “Bond formation and fiber embedment during ultrasonic consolidation,” Journal of Materials Processing Technology, 209, pp. 4915-4924.[13] Sun, Q., Rizvi, G. M., Bellehumeur, C. T., Gu, P., 2008, “Effect of processing conditionson the bonding quality of FDM,” Rapid Prototyping Journal, 14, pp.

 event-driven software architecture for process analysis in Additive ManufacturingMatthias Moi*, Christian Lindemann*, Ulrich Jahnke*, Rainer Koch*Chair of Computer Application and Integration in Design and Planning (C.I.K.) and Direct Manufacturing Research Center (DMRC), The University of Paderborn, Pohlweg 47-49, 33098 Paderborn, GermanyAbstractAdditive Manufacturing is still not commonly accepted as a considerable manufacturing process for serial products. Build rate and cost estimation or even the traceability of parts concerning Quality Management issues is weak. That's also because of the unavailability of adequate software solutions. Self-made solutions with Spreadsheets are often hard to adapt and inflexible in usage. This paper presents a distributed event-driven software architecture for cost assessment and traceability from powder to finished products.

This paper presents a distributed event-driven software architecture for cost assessment and traceability from powder to finished products. Further the approach of event-driven processing for cost calculation following the activity based costing methodology is discussed. The methodology considers arbitrary events (e.g. machine data or market prices) that may have an effect for a more detailed process analysis.IntroductionAdditive Manufacturing (AM) is a fast growing technology in several domains like the aerospace industry. On the first of June 2013 the FP7 project RepAIR started successfully. The consortium performs research on future repair and maintenance for the aerospace industry using AM. The development of novel IT solutions is also a significant part of the project. The simultaneously increasing complexity of processes where decisions may influence quality and costs of parts, directly and indirectly as well, the need of information technology becomes important.

The simultaneously increasing complexity of processes where decisions may influence quality and costs of parts, directly and indirectly as well, the need of information technology becomes important. The importance is also shown by the development of commercial solutions and products like Materialise Streamics™ (Dennis Vandenbussche 2012) or AutoFab. In order to specify costs more precisely, cost-drivers have to be identified. In (Lindemann et al. 2012) we presented a costing model, with focus on metal additive manufacturing (MAM). The methodology is focusing on the four main processes:1. Build Preparation 2. Manufacturing 3. Post processing 4. Quality controlFor these main processes we identified cost drivers like CAD-Preparation or machine preparation. Our approach mainly based on activity based costing, where more indirect costs are assigned into direct costs (Leitner 2007).

Our approach mainly based on activity based costing, where more indirect costs are assigned into direct costs (Leitner 2007). In every main process IT can support engineers, designers and decision makers during the whole processes (1 to 4). From the IT point of view heterogeneous data from several sources (e.g. machine parameters or requirements on the part) has to be digitalized or synchronized in order to increase the traceability of a part during the whole processes. The traceability is also important for quality management processes. Through continuous gathering of data a continuous improvement of quality management and quality assurance processes can be achieved. Furthermore collected data may influence cost drivers or quality drivers in each process and therefore in each of their sub-process.Based on these findings we are presenting an easy extendable software architecture that is based on the event-driven service oriented software architecture paradigm (cf. Mühl et al.

Mühl et al. 2006; Taylor 2009) and complex event processing (Etzion and Niblett 2011) to improve the traceability and simplify the cost estimation. Because of the fast development in additive manufacturing, a highly adaptable and configurable system with highly decoupled components is needed to keep in trying. The general implementation of the system as well as implementation details is not the subject of this work.Event Processing Even if the theory behind event processing is not new, the event-driven thinking enables us to compute complex issues by defining rules, usage of logical reasoning or pattern matching. An event may be everything that can happen in the real world. Etzion and Niblett defined this as follows:“An event is an occurrence within a particular system or domain; it is something that has happened, or is contemplated as havening happened in that domain.

The word event is also used to mean a programming entity that represents such an occurrence in a computing system” (Etzion and Niblett 2011).Let’s assume we would like to build a simple quality assurance alarm system for the observation of a build job. If there is an event type LowTemperatureDetected we can define a rule that sends an alarm to some system or tracks the information. Even this easy example shows how events can be processed. To get a little bit more complex, let’s assume that we have the additional event type BuildJobStarted. By defining a simple rule we are able to define what should happen, if those two events occur. An example could be to stop the build job (which is also an event) and inform a responsible person in order to avoid wasting material. This example is illustrated in Figure 1. For each sequence of events arbitrary rules can be defined. Each performed action may generate further events.

This example is illustrated in Figure 1. For each sequence of events arbitrary rules can be defined. Each performed action may generate further events. Therefore according to Etzion definition, everything that has happened can be interpreted as an event.2TIMELowTemeratureDetectedBuildJobStartedCancelBuildJobBuildJobCanceled4IF 1 AND 2 occured THEN cancel build job31Figure 1 Event processing exampleThe process of events generating further events is called complex event processing. Each sequence of events from an event stream can also be interpreted as a scenario. In the example above the scenario can be described as follows: Detection of build jobs which started with a low temperature in the building chamber.Event processing ApplicationsThe general (abstract) structure of an event-driven application is shown in Figure 2. The Event producers (Sensors, AM-Machines, Human-Interaction) are the event generators.

The Event producers (Sensors, AM-Machines, Human-Interaction) are the event generators. Theconsumers are applications, systems or components consuming these events (e.g. displaying them). The producers may be hardware sensors (e.g. heat detectors) or even software components. Note, that an event producer may also be an event consumer. The Event Processing block asynchronously interconnects producers with consumers. In many event- driven applications this is done by a messaging system like a message bus where components send and receive event objects from the message bus. An event object is an instantiated object of a specific event type.Event ProducerEvent ProducerEvent ConsumerEvent ProcessingEvent ConsumerFigure 2 Event ProcessingSummarizing (Luckham and Frasca 1998, S. 1) defined event processing as follows:“[...] event processing is a new technology for extracting information from distributed message-based systems.

1) defined event processing as follows:“[...] event processing is a new technology for extracting information from distributed message-based systems. This technology allows users of a system to specify the information that is of interest to them. It can be low level network processing data or high level enterprise management intelligence, depending upon the role and viewpoint of individual users. And it can be changed from moment to moment while the target system is in operation.”Event typesThere are many existing types of events. In a software system they are called event types that have to be defined for a domain.

In a software system they are called event types that have to be defined for a domain. Etzion and Niblett define event types as follows:“[…] An event type is a specification for a set of event objects that have the same semantic intent and same structure; every event object is considered to be an instance of an event type.” (Etzion and Niblett 2011)Transported information by an event object is specified by a set of attributes in its event type definition. An attribute is a key-value pair. Attributes may be required or not. Event type description attributes like the event type identifier are required for each instanced event object. Generally attributes can be distinguished as follows: Header attributes, payload attributes and open content. Header attributes are the system defined attributes, e.g. for the identification of the event object. These attributes are mostly generated automatically by a system. Payload attributes containing the data to be transported, resp.

After each step e.g. in the build preparation, new data gets available and can be used in arbitrary ways. The event-driven approach supports this thinking. Further the approach makes it easy to achieve a high decoupling between components. Dependency will only exist on required information (semantic level – inputs/outputs). E.g. components for the cost calculation in the post processing depend on the data of components supporting or observing the manufacturing process. Technically these components may work without their knowledge about the subsistence of each other. This makes the exchange of existing components as well as the adding simple. Hence an event processing network can be build. It exists of event producers, event consumers and event processing agents (Etzion and Niblett 2011).

Hence an event processing network can be build. It exists of event producers, event consumers and event processing agents (Etzion and Niblett 2011). The general idea of software agents relies on the definition of (Wooldridge 2002), were software agents are defined as independent computer systems interacting with their environment to achieve their goals. It also includes the communication with other agents. E.g. an event processing agent (EPA) may have the goal to store or process events, which are relevant to calculate the costs in the pre-processing, resp. the build preparation. EPA’s acting autonomously and can be replaced by other agents or new versions of agents without the modification of other agents. In case of the fast developments in the AM domain, this is a very important fact, because you are very flexible in the extension and adaption of software components, resp. EPA.Activity based costing with event-driven processing In (Lindemann et al.

EPA.Activity based costing with event-driven processing In (Lindemann et al. 2012) we analyzed product lifecycle costs for a better understanding of cost drivers in AM. Our approach is based on the activity based costing methodology (ABC) (cf. Miller and Vollman 1985; Cooper R. and Kaplan R. S. 1988) . The general requirement to model event processing agents (EPA) is to have access to relevant data in order to calculate fixed costs, variable costs and overhead costs. Hence we assume that EPA have access to databases were material and costs related information is stored. The results of the applicability are shown in Table 3. Since the general assumption of Luckham, that everything that has happened can be an event, the application of the ABC methodology in event processing systems can be done.Table 3 Applicability of event processing for ABCType of costs Fixed costs / Overhead costsVariable costsDescriptionCosts that are fixed in the whole process. E.g.

E.g. costs for machine preparation or labor cost rates. Costs for administration etc.Costs for material, energy consumption during the process, manufacturingApplicability of event processingThe EPA must get access to relevant data. EPA must be able to update costs rates if necessary, e.g. due to environmental changes or changes at the market. By adding a new order/build job, events are generated, received by the agent and processed. All relevant data to estimate and calculate fixed and the proportional overhead costs are stored in databases (even distributed databases) and can be accessed from an EPA for further processing. EPA must be able to observe the processes. In the build part preparation. Decisions (e.g. material selection, build part orientation etc.) aremachine cost rates.taken by the designer/engineer. Each decision may affect the resulting costs. While modeling event types (e.g.

material selection, build part orientation etc.) aremachine cost rates.taken by the designer/engineer. Each decision may affect the resulting costs. While modeling event types (e.g. machine or material selected), an EPA is able to process the data to pre- calculate the costs for the desired build job. Same holds for the manufacturing and the post processing. During and after the manufacturing events may influence the estimated costs.Event-Driven Architecture in AM In the previous sections we introduced how event-driven architectures can be designed with concept of event processing networks and event-driven agents. In order to support the whole process from powder to product we will now present the design of our event-driven architecture to achieve traceability and to calculate/estimate costs based on the ABC approach for the whole AM process.The general architecture is divided into five main subsystems (see Figure 5). A subsystem encapsulates functionalities logically.

A subsystem encapsulates functionalities logically. Beside the event-driven approach the architecture follows the Model-View-Controller pattern, where the view (application subsystem) is updated by the model (storage subsystem, Logging retrieval, message oriented middleware) through the controller (Agent Based Event Processing). The user interacts with the controllers indirectly.andusesCONTROLLERVIEWmanipulatesUSERMODELupdatesusesFigure 3 MVC patternIntegration SubsystemThis subsystem coordinates the whole event-driven communication between the subsystems and integrates external systems like AM machines to get process data during the manufacturing. The Message Oriented Middleware (MOM) provides a message bus where subsystems, resp. software components, can subscribe arbitrary event types. Published event objects are forwarded to components that have a subscription for associated event types.

software components, can subscribe arbitrary event types. Published event objects are forwarded to components that have a subscription for associated event types. The principle of MOM’s event bus is illustrated in Figure 4.Component 2Component 1Component 3Event busPublish E1Publish E2Subscribed for E1 and E2Figure 4 Event bus of the MOMThe communication via the MOM is asynchronous as by the definition of the messaging system via a message bus, resp. an event bus.Figure 5 Event-driven architectureApplication SubsystemThe application subsystem is composed of four main components; the Preparation, the Manufacturing, the Post Processing and the Quality Control. The structure is based on the four identified main processes to estimate the product lifecycle costs during the whole AM process. The components in the application subsystem are designed to support the processes, especially the designers and engineers from powder to product.

The components in the application subsystem are designed to support the processes, especially the designers and engineers from powder to product. The main tasks of the components are listed below.PreparationDefinition of requirements relevant for the manufacturing process  Selection of machine and material  Definition of quality assumptions o Roughness o Surface quality…Visualization of energy consumption regarding the parts to be build  Visualization of material consumption for support structuresManufacturingObserve build jobs  Show relevant events to the user  Visualize actual vs.

target performance by comparison of requirements and real-time data like machine data.Post-ProcessingObservation of post processing activities like heat-treatmentVisualization of estimated costs for post processingQuality controlVisualization of costs for quality control regarding specified requirements  Documentation of new knowledge regarding quality controlThe components in the application subsystem are representing the view, resp. the interface to the user. They are designed to consume events for information visualization for the user and to produce events by the interaction with the user. While interacting with the view, events are generated for further processing in the controller (Agent Based Event Processing subsystem).Agent Based Event ProcessingThe Agent Based Event Processing subsystem covers the intelligence of the whole system. For each process specialized event processing agents (EPA) are defined.

For each process specialized event processing agents (EPA) are defined. While components in the application subsystem are event producers and event consumers, EPA processing the events. This includes the generation of new, complex events by defined rules and pattern matching mechanisms. The main tasks of the EPA are listed below.PreparationCost Estimation in the preparation  Calculation of estimated material and energy consumption regarding the specification done by the users and machine selection.Calculation of build time and time for post processing to rearrange parts in the building chamber.Estimation of energy consumption regarding the parts to be build  Estimation of material consumption for support structuresManufacturingObserve build jobs  Processing of relevant events during the manufacturing process like critical situations recognized by defined rules or pattern matching.Visualize actual vs.

target performance by comparison of requirements and real-time data like machine data.Post-ProcessingCalculating of costs for post processing activities like heat-treatment  Estimation of post processing costs  Estimation of costs for support structure removal.Quality controlEstimation of costs for quality control and improvement of quality assurance methodologies regarding specified requirements.  Analysis for process improvement by collected data.  Correlation analysis between different build jobs for quality improvement and costreduction.Storage Subsystem and Logging/RetrievalThe storage subsystem covers databases (or even knowledge bases) to process and track all activates in the system. It builds the model of the whole system. The Storage subsystem provides an interface where components of the other subsystem can request and store data.

It builds the model of the whole system. The Storage subsystem provides an interface where components of the other subsystem can request and store data. The possible huge amount of data makes it possible to make further analysis like correlation analysis between chosen parameters in the build preparation.The Logging and Retrieval subsystem is an event consumer which is subscribed for all event types at the MOM. This guaranties maximum event based traceability, because each event is stored. Event objects stored in the database can be reproduced for arbitrarily time frames..Conclusion and future work In this paper we’ve introduced the paradigm of event-driven architectures and showed the potential of application in the additive manufacturing domain. We’ve analyzed the applicability of event-driven processing for the activity based costing methodology.

We’ve analyzed the applicability of event-driven processing for the activity based costing methodology. Based on these findings we designed an event-driven architecture to support the whole AM process in order to achieve a high traceability and an easier calculation of product-lifecycle costs by the usage of the event processing agents approach.Regarding the architecture the next main steps will include be the implementation of the architecture based on existing software frameworks like ESPER 1, Oracle Complex Event Processing2 or HornetQ3.Further investigations will be made in the identification and specification of process relevant event types and therefore the definition of event processing agents. Since everything that has been happed can be an event the challenge becomes which occurrences matter. Many occurrences throughout AM process are very process-dependent, which means that deep process analyses have to be made.

Many occurrences throughout AM process are very process-dependent, which means that deep process analyses have to be made. The relevance of event types differs from process to process. E.g. usage of different materials may require different post processing steps were different events may occur or be relevant. Therefore we will focus on metal selective laser sintering processes in the initial phase.This comes with research in rules and pattern for event-driven cost estimation of product- lifecycle costs in additive manufacturing. The aim is to define event-driven agents in order to derivate event-driven costing methodologies of existing activity based costing approaches.1 http://esper.codehaus.org/ 2 http://docs.oracle.com/cd/E13157_01/wlevs/docs30/pdf/get_started.pdf 3 http://www.jboss.org/hornetqLiterature Cooper R.; Kaplan R. S.

AMINATION OF BUILD HEIGHT IN ULTRASONIC CONSOLIDATION FOR FOIL WIDTH SPECIMENS USING SUPPORTSJames M. Gibert1, 3, Daniel T. McCullough2, Georges M. Fadel3 Kenneth E. Johnson4 2Corvid Technologies 3Department of Mechanical Engineering, Clemson University 4Solidica, Inc.ABSTRACTUltrasonic consolidation (UC) is a novel, solid-state, additive manufacturingfabrication process. It consists of ultrasonic joining of thin metal foils and contour millingto directly produce functional components in a variety of geometries. The bond betweenlayers forms when an ultrasonic horn creates a local oscillating stress field at the matingsurfaces. It is commonly theorized that the high frequency vibration under pressureproduces a metallurgical bond without melting the base material.

It is commonly theorized that the high frequency vibration under pressureproduces a metallurgical bond without melting the base material. The mechanism behindthe bond is believed to be due to interfacial motion and friction that disrupts surfacecontaminants, arguably allowing direct metal to metal contact, and producing sufficientstress to induce plastic flow and promote the growth of grains across the mating surfaces.Ignored in this explanation is the role of substrate dimensions on the quality and strengthof the joining process. Researchers have previously examined the effective heightlimitations of the build process, i.e., the limiting height to width ratio of one of thecomponent features being fabricated. This paper extends the experimental work on usingsupport materials to extend build height on specimens using two different candidate1 Corresponding author385materials, tin bismuth, and a mixture of sugar, corn syrup, and water, referred to as“candy”.

Tin bismuth and candy the represent the extremes of a tradeoff betweenconvenience and stiffness that a support material must possess.INTRODUCTIONUltrasonic consolidation (UC) also known as Ultrasonic Additive Manufacturingis an innovative, solid-state, rapid manufacturing fabrication process. Solidica, Inc.originally developed the process, which is composed of ultrasonic joining of thin metalfoils and contour milling to directly produce functional components. The process hasbeen utilized in applications ranging from embedding electronics into armored vehicles,embedding shape memory alloys within structures, fabricating injection mold tooling andfabricating lightweight structural panels for satellites [1].DESCRIPTION OF ULTRASONIC CONSOLIDATIONThe UC process begins with a thin metal foil being placed on a sacrificial baseplate.

The base plate is bolted downward and heated to 300°F (approximately 150°C).The ultrasonic horn compresses the foil while simultaneously vibrating transversally at anominal frequency of 20 kHz and at amplitudes ranging from 1.97x10-1 to 1.18 mils (5-30 µm) while simultaneously rolling over the foil. The foil is typically 5.90 mil (150 µm)thick and 0.94 inches (23.88 mm) wide. After the foil is bonded, the process is repeatedfor additional foils across the width of the build plane for each layer of the desired endcomponent. The consolidated foils are machined as needed throughout the build processto produce the desired final part geometry.

The consolidated foils are machined as needed throughout the build processto produce the desired final part geometry. Figure 1 illustrates the process.38612 Sonotrode presses on foil while rolling and vibra!ng at 20 khz with an ampitude of 16μm.Surfaces pushed together by compressive load.Thin Metal FoilsBuild plate3Fric!on at interface breaks up oxides.4Mechanical interlocking, plas!city, grain growth.Figure 1: Overview of the UC processPREVIOUS RESEARCH ON BUILD HEIGHT LIMITATIONRobinson et al. [2] found that an apparent limit on build dimensions existsthrough a series of experiment where they varied the width and orientation of thespecimen with respect to the vibrating sonotrode and consolidated layers until failureoccurred. Specifically, they examined five specimens at the following widths: 0.94inches, 0.5 inches, 0.25 inches, 0.125 inches and 0.063 inches. Each specimen has alength-to-width ratio of 10:1.

Specifically, they examined five specimens at the following widths: 0.94inches, 0.5 inches, 0.25 inches, 0.125 inches and 0.063 inches. Each specimen has alength-to-width ratio of 10:1. The failure of the smaller width specimens was attributed toinsufficient bonding area. Conversely, for specimens of larger widths they concludedthat layers couldn’t be bonded to a feature if its dimensions reach a critical value,specifically, if its height-to-width ratio (h/w) is approximately 0.7 to 1.2, hereafterreferred to as high aspect ratio features. They attributed the inability to bond, to thesubstrate becoming increasingly compliant, i.e., lowering its stiffness and reducing theamount of differential motion between the specimen and bonding tape. This assertionsupports anecdotal evidence by Solidica, Inc. that support material added to high aspectratio (height-to-width) features, extends the height to which the feature can be built and387bonded successfully.

that support material added to high aspectratio (height-to-width) features, extends the height to which the feature can be built and387bonded successfully. Solidica demonstrated that utilizing a support material couldachieve high aspect ratio features of over 20 (figure 2).Figure 2: High aspect ratio feature produced with the use of support materialZhang and Li [3] attempted to explain the findings of Robinson et al. [2] throughnumerical simulations. The researchers performed a fully transient, dynamic, finiteelement analysis that allowed for elastic, plastic and thermal strain. The simulationshowed that, as h/w approaches unity and beyond, the magnitude of the frictional stressdecreases, and that concentrated bands of high shear strain appear in the substrate as h/wapproached one. They attributed this to positive wave interference that resulted inminimum interfacial displacement.

They attributed this to positive wave interference that resulted inminimum interfacial displacement. The researchers later refined their model [4] to clarifyhow the waveform causes bonding degradation through a two-dimensional model of aquarter domain of the substrate. The researchers modeled the wave motion using aRayleigh Ritz analysis. They argue that in steady state shear; stress concentrations appearon the substrates edge. After half a vibration cycle the concentrations switch locationsresulting in wave interference and in minimal shearing stress due to the two waves388interfering. The interference results in insufficient elastic plastic deformation to promotebonding. Relating both the wave interference model and the ratio of energy transferredinto the bonding area to the frictional energy, they conclude that an h/w equaling oneresults in a limit on bonding energy.

However, it is unclear from their work why thisgeometry results in wave interference.While Zhang and Li [4] used a Rayleigh-Ritz model, the first use of this techniqueto understand the UC process was by Gibert et al. [5] . They showed that for featuresbuilt at the nominal tape width with high aspect ratios, several modal frequencies of thefeature approach the 20 kHz excitation frequency of the sonotrode. They also showed thatthese modes are only weakly dependent upon the length of the build piece. Theypostulated that the increase in compliance of the build feature could be due to anexcitation of a modal frequency by the sonotrode. In a follow up study [6] the researchersexperimentally verified their hypothesis for nominal tape width specimens.

In a follow up study [6] the researchersexperimentally verified their hypothesis for nominal tape width specimens. Using bothlumped parameter and finite element modes of the UC process they accurately predictedtrends in vibration that can be correlated to the build height limit at nominal tape widthsbased on examining differential motion at the interface. Both models indicated that oncethe region of experimentally observed regions of bond degradation due to height arepassed it is possible to re-initiate bonding. Experimentally, they verified the models’predictions by consolidating a large stack; milling it down to a high aspect ratio featureand then resuming the consolidation process, see Figure 3. In this manuscript we expandon the tests, conducted by Gibert et al.

In this manuscript we expandon the tests, conducted by Gibert et al. [6], by testing the effectiveness of supportmaterials in extending the limiting build height in the UC process.389Figure 3: Welding of additional layers on specimen of 2.00 X 0.94 X 2.50 inches (50.80 X 23.88 X 63.50 mm) [8]PREVIOUS RESEARCH ON THE USE OF SUPPORT MATERIALSThe use of support material is not original in Ultrasonic Consolidation; their usewas motivated by anecdotal evidence by Solidica, Inc. indicating that the addition ofsupports increases the build height of freestanding features.

indicating that the addition ofsupports increases the build height of freestanding features. Additional work conductedby Solidica with support from the US Department of Defense and the National Center forManufacturing Sciences in 2005 and 2006 established the earliest methods for theautomated delivery of a tin-bismuth support material within an Ultrasonic Consolidationplatform.390The first published use of support materials in UC that included an investigationof feature building effects was by Swank and Stucker [7]. They examined the effect ofsupport materials using three build configurations: an enclosed pocket, freestanding rib,and an open channel. In addition, they used five candidate support materials: metal alloy(tin-bismuth), a thermoplastic (Water WorksTM), a thermoset (Leco quick cure (QC)epoxy), a wax (water soluble casting wax), and an organic (aluminum filled sucrose).They found that only tin-bismuth allowed all geometries to be built.

In addition, theyproposed several guidelines for the use of support materials. The recommended that oneuse a lower build temperature to reduce softening of the support material. They noted thatharder support materials reduce the amount of interface voids and delamination, andmetal support materials are preferable to other materials in that they allow depositedlayers to be welded to their surface. Metals allow subsequent layer bonding on thesupport material. They also found that localized heating during the UC process can meltsupport materials. However, left unanswered is the exact improvement of supportmaterials on extending the build height in the process.This paper examines the effect of support materials on the maximum build heightobtainable for the 0.94-inch width specimens. In our work, we differ in both supportgeometry and material than those used by Swank and Stucker [7]. They used enclosedcavities.

In our work, we differ in both supportgeometry and material than those used by Swank and Stucker [7]. They used enclosedcavities. In this study, we use semi-enclosed support materials, i.e., two faces of thematerial are free, and the support material is not flush with the build specimen’s height.We considered this geometry for several reasons. First, this arrangement may be seen asthe limiting configuration of applying the support material that would allow it to stiffenthe build specimen. Second, it is easily foreseeable that during the use of UC one may391want to produce components where the enclosed cavity would be problematic. Thirdly,by varying height we prevent localized melting of support materials that can occur if thematerial is directly below the sonotrode. Finally, this geometry is of particular interestconsidering the work of Gibert et al.

Finally, this geometry is of particular interestconsidering the work of Gibert et al. [6]; it adds only side supports to free standingfeatures.ROLE OF SUPPORT MATERIALIn simple mechanical terms, the stiffness of the build piece can be represented bysprings, Figure 4, through which forces are transmitted to the ground. The supportmaterial can be viewed as a spring in parallel to build feature stiffness.Figure 4: Conceptual model of the effect of support materialThe addition of support materials has two possible effects on a free standing buildfeature: 1) changes the lateral stiffness of the build feature resulting in changes of modalfrequencies of vibration of the build feature if the support material adheres to sides of thefeature, and 2) increases dampening in the structure. However, the ideal support material392must do more than change the structural characteristics of the build piece.

However, the ideal support material392must do more than change the structural characteristics of the build piece. It mustconvenient to remove when the consolidation process is complete. Figure 5 shows ahypothetical tradeoff between stiffness and convenience. In our studies, we examined thetwo extremes of the tradeoff, SnBi, and a mixture of hardened sugar, water and cornsyrup we call “candy”. Material characterization shows that they have modulus of 60.8GPA and 16.14 GPA, for SnBi and candy, and Brinell Hardnesses of 3, and 15,respectively [8].Figure 5: Tradeoff between stiffness of support material and ease of removal393Figure 6: Specimens before filling used in support material characterizationBUILD SPECIMENS MATERIAL AND GEOMETRYThe build specimens were constructed from aluminum foil 3003-H18 provided bySolidica, Inc.

The dimensions of the build plates were 14.00 x 14.00 x 0.51 inches (356 x356 x 13 mm) thick 3003 H-18 plates, again provided by Solidica, Inc. The specimenswere milled to a height, width and length of 0.94 x 0.94 x 2.50 inches, i.e., at theproblematic aspect ratio for nominal tape width specimens.SUPPORT MATERIAL COMPOSITION, GEOMETRY AND SPECIMEN PREPERATIONWe chose three geometries for the support material: 0.25 x 0.25 x 2.50 inches,0.50 x 0.50 x 2.50 inches, and 0.75 x 2.00 x 2.50 inches. This was a preliminary studyand these geometries would give us two points to quantify the effect of support height onbuild limit and one extreme point in examining the effect of an oversized support. We394conducted the test for all the geometric configurations using both SnBi and the candysupport material totaling six tests. The SnBi ingots were heated in a furnace to melting,roughly 450 °F. The candy was made of a mixture of sugar, corn syrup, and water.

The SnBi ingots were heated in a furnace to melting,roughly 450 °F. The candy was made of a mixture of sugar, corn syrup, and water. Interms of mass the candy consisted of 82 % percent household sugar, 15% corn syrup, and3% water. The mixture heated to roughly 300oF for 20 minute. Each support materialwas poured in its molten state and bracing material was used to ensure the edge of thesupports remained flush with both the build specimen and the support bracing. Thecandy material proved to be somewhat unwieldy and the material overfilled the trough.We were able to ensure that SnBi was flush with the bracing. In each case the supporttrough was totally filled. Figure 7 and 8 show the filled specimens before welding.

In each case the supporttrough was totally filled. Figure 7 and 8 show the filled specimens before welding. Afterthe specimens were placed on the machine, we waited two hours before beginning thetest to allow the specimens to reach thermal equilibrium.395Figure 7: SnBi used in support material characterizationFigure 8: Candy used in support material characterization396PROCEDURETable 1 presents the geometry and UC process parameters for the tests. The baseplate was heated to 120 °F (50 °C) for the candy and 200 °F (93 °C) for SnBi. The lowerbase plate temperature was used to avoid exceeding the melting point for both materialswhich would occur at the standard operating temperature 300 °F (150 °C) [7]. Inaddition, the normal load had to be lowered for each support material. Initially, thenormal load was set to 1400 N (314.7 lbs.).

Inaddition, the normal load had to be lowered for each support material. Initially, thenormal load was set to 1400 N (314.7 lbs.). This proved to be problematic since thepower supply would short out and additional layers could not be welded over thespecimen; the same behavior was observed when welding over the freestandingspecimens of this dimension by Gibert et al. [6]. We systematically reduced the normalload until bonding occurred at a load of 269.8 lbs. (1200 N) and 179.8 lbs. (800 N), forthe SnBi and candy, respectively. The amplitude was set to 9.84x10-4 inches (25 µm) andthe rolling speed of 100 in/min (42.33 mm/s).

(1200 N) and 179.8 lbs. (800 N), forthe SnBi and candy, respectively. The amplitude was set to 9.84x10-4 inches (25 µm) andthe rolling speed of 100 in/min (42.33 mm/s). Each specimen was then welded for up to55 layers or until bond failure was observed.397Table 1: Test Specimens and process parameters used in support material test with a rolling speed of 100 in/min (42.33 mm/s) and amplitude of 9.84x10-4 inches (25 µm)CFGSupport HeightSupport WidthSupport MaterialBase Plate Temp.Normal Load10.25 in0.5 inCandy120 °F179.8 lbs. (800N)20.50 in0.5 in(50 °C)30.75 in2.0 in40.25 in0.5 inSnBi200 °F269.8 lbs. (1200N)50.50 in0.5 in(93 °C)60.75 in2.0 inRESULTS OF SUPPORT MATERIAL TESTSTable 2 presents the results of the support material test.

(1200N)50.50 in0.5 in(93 °C)60.75 in2.0 inRESULTS OF SUPPORT MATERIAL TESTSTable 2 presents the results of the support material test. Although the use of thecandy required a drastic reduction in normal load, its performance was superior to SnBi.In two of the three configurations, specifically, 1 and 3, the candy showed no signs offailure at 55 layers. Conversely, only one specimen of SnBi, configuration 5, showed nosign of failure. The candy was observed to progressively fracture as successive layerswere welded, Figure 9.398Figure 9: Cracking of candy support materialThis was expected due to the much greater brittleness of the candy compared to therelatively soft and ductile SnBi. Despite cracking, the candy specimens remainedattached to the build feature for the duration of the UC process. In contrast, the SnBispecimens showed no signs of cracking in their bulk.

Despite cracking, the candy specimens remainedattached to the build feature for the duration of the UC process. In contrast, the SnBispecimens showed no signs of cracking in their bulk. Figure 10 shows that supportstructures did not remain “wedged” between the specimen and the support bracing butwere vibrated loose and shifted during the weld process. Note that the height or width ofthe support material was not a factor in the performance of the candy.

Surprisingly, thetallest and widest SnBi sample refused to bond.399Table 2: Test results of observed failure for candy (sucrose mixture) and tin bismuth (SnBi) support materialsCFG Support HeightSupport WidthMethod of FailureLayer at FailureTotal Height10.25 in0.50 inNo Failure551.27 in20.50 in0.50 inCrack Formed481.22 in30.75 in2.00 inNo Failure551.27 in40.25 in0.50 inCrack Formed281.11 in50.50 in0.50 inNo Failure551.27 in60.75 in2.00 inDetachment of 1st layer241.08 inFigure 10: Displaced SnBi support material400Support MaterialCandySnBiDISCUSSION OF SUPPORT MATERIAL RESULTSWe hypothesize that the superior performance of the candy is primarily due totwo factors: 1) propensity oft the candy to wet to the aluminum, and 2) a mismatch in thethermal properties between SnBi and aluminum that may have left a gap after initialheating cooling and reheating.

These two factors outweigh the relative hardness orstiffness of the materials; thus, SnBi’s stiffening effect could not be fully realized.ADHESION OF SUPPORT MATERIALS TO ALUMINUMThe wetting depends on the relative surface energies of the molten liquidsupport material and the aluminum build specimen. When the surface energy of theliquid is greater than that of the solid; the liquid will have a large contact angle andappear as a bead. The surface energy also known as surface tension is a function of manyvariables including surface roughness and temperature; thus, obtaining the precise valuesfor the surface energy of the material would be difficult to obtain. Lee et al. [9] showsthat the equivalent surface tension of the molten SnBi is approximately 340 dynes/ cmmuch larger than the surface tension of the solid aluminum which is 35-45 dynes/cmdepending on the alloy.

The best estimates of the surface tension of the sucrose mixtureare near 65 dynes at room temperature [10]. While the authors did not specify the size ofthe capillary tube used to perform the test. Estimating the tube width to be roughly 2 to 4cm gives surface tensions of 16.25 dynes/cm to 32.5 dynes/cm very close to the surfaceenergy of aluminum.401THERMAL PROPERTIES OF SnBiCandy’s superior wetting ability may not be the only the reason behind itsperformance. The coefficients of thermal conductivity and thermal expansion of thematerial also play a significant role in how well each support the build specimens. Thisrelationship can be explained by considering the following simplified problem.

Thisrelationship can be explained by considering the following simplified problem. Assumethe heat flow and the change in length lie primarily in one-dimension, then usingFourier’s law and the simple definition of one-dimensional thermal expansion, the changein length due to heating isΔL =αL2Q κA,where α, κ, L, A, Q , are the coefficients of thermal expansion, conductivity, length indirection of heat flow, the cross-sectional area, and the heat flowing in the specimen.Fixing the geometry and input heat, the ratio of thermal conductivity to thermalexpansion determines the change in length. Table 3 presents the thermal properties usedfor aluminum and SnBi. The ratio of α κ for the aluminum and SnBi are 8.19e5 cm and1.07e7 cm, respectively. Indicating that a fixed length of SnBi will expand more thanaluminum, since it will absorb more heat. If the materials reach the same temperatureand are allowed to cool, the SnBi will shrink more.

If the materials reach the same temperatureand are allowed to cool, the SnBi will shrink more. This problem is complicated bydifferences in geometry between the supports and build feature, and that the heat flows inthree dimensions in the actual tests. We theorize that once the SnBi solidifies the supportstructure shrinks so that it is no longer in contact with the build specimen.402(1)Both materials are poured into the supports in their molten states, then allowed tocool to room temperature, and then re-heated to reduced operating temperatures of theUC process. As the material changes the temperature, the volume subsequently increasesand decreases.

As the material changes the temperature, the volume subsequently increasesand decreases. A thermal-mechanical analysis using the ABAQUS finite elementpackage was used to perform a rudimentary test of this hypothesis.Table 3: Thermal properties of aluminum, SnBi used in finite element simulationPropertiesAluminumSnBiThermal Conductivity Coefficient of Thermal Expansion19 W °C/cm 2.32e-05 /°C160 W °C/cm 1.50e-05 /°CIn the analysis, convective heat loss is neglected, SnBi is a assigned an initial temperatureof 450 °F, the initial molten temperature when poured into the trough, and it is modeledas totally filling the trough between the aluminum build specimen and support bracing,thus it occupies the same volume as the trough. A subsequent steady state analysis ofcooling is performed at an ambient temperature of 68 °F. Next we model the reheating inthe UC process by applying a 200 °F thermal boundary condition to the bottom of thebase plate.

Next we model the reheating inthe UC process by applying a 200 °F thermal boundary condition to the bottom of thebase plate. Figure 11 shows the resulting geometry changes resulting from this analysis.The support is 34 µm smaller than the width of the trough. This is a substantial differenceconsidering the peak-to-peak amplitude of the sonotrode during the tests was only 25 µm.The results obtained from this analysis are preliminary and cannot be taken as definitive.We use estimates of the thermal and mechanical properties of SnBi. The model neglectslatent heat, potential non-linear expansion and contraction, and temperature dependentmaterial properties. Furthermore, the SnBi undergoes a phase change when cooling from403its molten state to solid state; changing its mass density.

Furthermore, the SnBi undergoes a phase change when cooling from403its molten state to solid state; changing its mass density. The density change should serveto amplify the change in volume.Figure 11: Deformation in heated SnBi support material and base plate from FEA modelCRACKING OF CANDY SUPPORTFinally, the brittleness and resultant cracking of the candy suggested that, even inreheat during the UC process, the material was below its glass transition temperature. Theglass transition temperature of the mixture of sugar, glucose and water used in this studyis not readily available. Glass transition temperatures of similar substances range from140 °F (60 °C) to 163 °F (73 °C) [11] which is above the operating temperature of 120 °F(95 °C) used in these experiments.

Glass transition temperatures of similar substances range from140 °F (60 °C) to 163 °F (73 °C) [11] which is above the operating temperature of 120 °F(95 °C) used in these experiments. Interestingly, the cracking does not affect candy’s rolein extending build height.404EFFECT OF GEOMETRYThe results of the study and the limited sample size make it difficult to determinethe effects of the dimensions of the support material on the build limit. It is readilyapparent that the support material does not have to equal the height of the build specimenfor bonding to occur. The current set of experiments does not provide enough data pointsto predict the effect of width support material on build height. In the candy, we seesmaller height support of 0.25 inches allowed the specimen to weld to 55 layers withoutany signs of failure compared to the support of height of 0.50 inches in which thespecimen showed signs of cracking at 48 layers.

The exact opposite was seen in SnBiwith the crack formed when welding using a support of height of 0.25 inches. Finally,unaccounted for in this experiment is the effect of the dimension of support bracingwhich may play a prominent role in the support’s effectiveness.Figure 12: Summary support material properties on build height405RECOMMENDATIONSFigure 12 summarizes the effect of support material properties on extending bondheight. These results can be added to the original recommendations of Swank and Stucker[7] on the use of support materials. In using support material the surface adhesion andthermal expansion should be considered whenever possible. The surface energy of themolten support material and the build specimen should be as close as possible. Thematerials should have similar coefficients of thermal expansion when the supportmaterial is not completely enclosed.

Thematerials should have similar coefficients of thermal expansion when the supportmaterial is not completely enclosed. Finally, when using material with significantlydiffering coefficients of thermal expansion one may need to apply the support material inan enclosed channel.FUTURE WORKThis work can lead to several avenues of future research. First, the results herecan form the basis of an expanded support material study that considers the specimen andsupport material’s, height, width and geometry, as well as back support height, andgeometric configuration in extending build height using a design of experimentsapproach. Second, this study only considered two support materials. Theexperimentation should be repeated for the other materials such as those examined bySwank and Stucker [7].

Second, this study only considered two support materials. Theexperimentation should be repeated for the other materials such as those examined bySwank and Stucker [7]. This work illustrates that the design space of candidate supportmaterials needs to be expanded to include the wetting and thermal properties of thematerials.406ACKNOWLEDGEMENTSThis research was supported by a STTR grant from the National Science Foundationthrough the Solidica, Inc. The views presented in this paper do not necessarily reflectthose of our sponsor whose support we gratefully acknowledged. Thanks to Solidica,Inc. for their continued support throughout this research effort.REFERENCES1.2.3.4.5.6.7.8.9.10.George, J. and B. Stucker, Fabrication of Lightweight Structural Panels through Ultrasonic Consolidation. Virtual and Physical Prototyping, 2006. 1(4): p. 227- 241.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEXAMINATION OF THE CONNECTION BETWEEN SELECTIVE LASER-MELTED COMPONENTS MADE OF 316L STEEL POWDER ON CONVENTIONALLY FABRICATED BASE BODIESMartin Link1, Tobias Haefele1, Eberhard Abele11 Institute of Production Management, Technology, and Machine Tools (PTW), Technische Universität Darmstadt, Darmstadt, GermanyAbstractThe advantages of selective laser melting lie in the production of complex, small components in small batches. For large-volume components, the use of additive manufacturing (AM) processes is limited by the available installation space, low build rates, and high material costs. For the production of large and less complex workpieces, conventional manufacturing processes such as milling are more economical. The background of this study was to combine both processes to decrease manufacturing times.

The background of this study was to combine both processes to decrease manufacturing times. For this purpose, a body made of 316L (1.4404) steel powder was printed using selective laser melting on conventionally manufactured stainless-steel base bodies. The use of multi-materials enables optimized machinability in the respective manufacturing process. This paper examines the hardness properties of multi-material samples and uses micrographs to analyze the microstructure of their connection area. A complete connection between hybrid components made of comparable materials was determined.1. IntroductionAdditive manufacturing (AM), an emerging field in manufacturing technologies, features the principle of building solid parts directly from 3D-CAD data by the layer-by-layer addition of materials, which may be plastic, metal, concrete, or, as expected in the future, human tissue.

According to Wohler’s Report, 135 manufacturers produced AM systems in 2017 compared to 65 manufacturers in 2015. Today, established manufacturers are under pressure due to the huge number of new machines and unprecedented levels of competition in the additive production marketplace. In 2017, the AM segment of worldwide sales are increased by 20.9% to US $7.336 billion [1]. Powder-bed-fusion-based AM processes use thermal energy to selectively fuse regions of a powder bed [2]. This study focuses on the laser melting of metal powders, often also referred to as selective laser melting (SLM). Given the current state of the art, low construction rates and high material costs limit the economics of the SLM process for large-volume components [3]. In this study, the relationship between an additively built volume and a conventionally manufactured base body was investigated. The advantages of the SLM process lie in the production of complex components in small batches.

The advantages of the SLM process lie in the production of complex components in small batches. Due to their high rates of cooling, the components achieve increased strength values compared to those of cast components [4]. Conventional manufacturing processes, in contrast, are better suited for large quantities and less complex components. By combining additive construction using the SLM process with conventionally manufactured base bodies, the advantages of both processes can be obtained. In this paper, this combination of processes is referred to as a hybrid process. This hybrid process is already being used in industry. For example, large components are being produced by fine forging a base body and then using SLM to add complex structures. To be able to produce hybrid components for industrial use, multiple factors must be considered. The base bodies must have a flat surface of low surface roughness, which is403fixed parallel to the powder bed.

The base bodies must have a flat surface of low surface roughness, which is403fixed parallel to the powder bed. In the hybrid production of components, the connection between the two materials is of particular concern and must be examined in detail. In this regard, a material combination must be selected that enables effective hybrid production. Furthermore, the first powder layer must be of a defined height to enable process reproducibility. Tensile tests of hybrid specimens show that the tensile rods fail not in the material transition but in the weaker area of the two materials [5–7]. Furthermore, hybrid samples with different aspect ratios (1:2, 1:1, 2:1) of the hybrid samples were examined, which also showed no influence on the tensile strength [8]. Another aspect for consideration is surface roughness, which is a decisive factor in many applications. The fatigue performance of AM parts is primarily dependent on their surface roughness [9, 10].

Another aspect for consideration is surface roughness, which is a decisive factor in many applications. The fatigue performance of AM parts is primarily dependent on their surface roughness [9, 10]. To meet the requirements of hybrid components, the additive volume must be post processed, e.g., grinded, shot-peened, or machined.This paper investigates the connection between the base body and the additively applied volume. To do so, structural and hardness investigations were conducted in the connection area to identify critical factors in the industrial production of components in this area of investigation.2. Research methodologyThe experiments were conducted using a commercially available SLM system EOS M290 from EOS GmbH (Germany), which uses a 400 W Yb fiber laser focused on a 100-μm spot. In all the experiments in this study, a gas atomized stainless steel 316L powder with a particle distribution range of 11 µm to 54 µm and a d50 of 25 µm was used.

In all the experiments in this study, a gas atomized stainless steel 316L powder with a particle distribution range of 11 µm to 54 µm and a d50 of 25 µm was used. The 316L_SurfaceM291 parameter set qualified by the machine manufacturer was used for the additive build-up and the layer thickness was set to 20 μm. The conventionally manufactured base bodies consist of 1.4301 stainless steel. Tables 1 and 2 list the chemical compositions of the 1.4301 and 1.4404 steel materials.Table 1: Chemical composition of the 1.4404 steel material [11]Element FeCrNiMo C Mn CuPSSiMin. Max.Balance 17.00 19.0013.00 15.002.25 3.000.032.000.500.0250.010.75Table 2: Chemical composition of the 1.4301 steel material [12]ElementFeCrNiCMnPSSiMin. Max.Balance17.50 19.508.00 10.500.072.000.0450.0151.00The 252 mm × 252 mm × 20-mm substrate plate has a grid of through holes.

Max.Balance17.50 19.508.00 10.500.072.000.0450.0151.00The 252 mm × 252 mm × 20-mm substrate plate has a grid of through holes. The base-body measures 10 mm × 10 mm × 25 mm and has 13-mm deep M4 threaded holes. The additively built body measures 9 mm × 9 mm × 10 mm. 20 base bodies were fixed to the substrate plate with Allen screws (Figure 1).404N0.1N0.111.44041.4301Figure 1: Schematic illustration of the hybrid samples on the substrate plate with base bodies of 1.4301 and additive built- up volumes of 1.4404The surface plane of the base bodies is ground to ensure a homogeneous roughness profile and a parallel plane between the specimens and the building platform. After the building process, the components were cleaned in an ultrasonic bath and the samples embedded with Struers EpoFix and EpoFix hardener (ratio 25:3), which then hardened for 15 hours.

After the building process, the components were cleaned in an ultrasonic bath and the samples embedded with Struers EpoFix and EpoFix hardener (ratio 25:3), which then hardened for 15 hours. The surfaces of the hybrid specimens to be examined are ground in stages using 180-, 320-, 500-, 1000-, 2500- and 4000-grit abrasive paper. In the final polishing process, a surface roughness of Rz 1 µm was achieved. Hardness tests were performed with Struers' DuroScan20 semi-automatic hardness testing system and Vickers low-load and microhardness tests.The indenter used in Vickers hardness measurement consists of a square diamond pyramid with a square base whose sides are inclined 136° to each other. Due to the proportionality of the test load to the indentation surface, Vickers hardness values are load-dependent. For particularly low loads, however, properties of the material, e.g., its surface tension, have an effect on the test load.

For particularly low loads, however, properties of the material, e.g., its surface tension, have an effect on the test load. For this reason, in addition to the microhardness tests, test loads with low-load hardness were performed for comparison. 15 HV0.05 measuring points were used for microhardness measurement and three measuring points in the connection area for low- load hardness measurement. The Vickers hardness (HV) values are calculated by (Eq.1), as follows:𝐻𝑉 =(0,102 ∗ 𝐹) 𝐴= (2 ∗ sin (68°) 𝑑) ∗ 0,102 ∗ 𝐹(Eq.1)where F is the test force, A is the area, and d is the mean value of the diagonals. The holding time is 7 seconds [13]. To make the structures of the hybrid samples visible, immersion etching was performed in a Behara 2 color etching mixture.

The holding time is 7 seconds [13]. To make the structures of the hybrid samples visible, immersion etching was performed in a Behara 2 color etching mixture. This is a precipitation estimation in which the cut surface reacts with the etchant, whereby the thickness and color of the precipitation layer is influenced by the crystal orientation and phase structure of the material. To obtain an overview of the structure, an optical microscope LEICA DM6000 was used for the examinations. To analyze the heat-affected zone, we used energy dispersive X-ray spectroscopy.With a scanning electron microscope from Phenom-World BV 15 measurements were performed at a distance of 100 µm from the edge of the hybrid part. The acceleration voltage used was 15 kV.To investigate the heat- affected zone, a single powder layer was additively built up and examined on additional samples.4053.

The acceleration voltage used was 15 kV.To investigate the heat- affected zone, a single powder layer was additively built up and examined on additional samples.4053. ResultsThe macroscopic examination of the hybrid samples revealed no cracks or discolorationof the material in the connection area.The surface roughnesses of the two areas were quite different, as shown in the inset of Figure 2.Figure 2: Produced hybrid samples on the substrate plateThe investigations with optical and scanning electron microscopes showed neither cracks nor larger visible porosities in the connection area of the hybrid samples. This connection is characterized by melting traces from the laser radiation. The widths and depths of the melt tracks vary. The area of the base body 1.4301 shows strong line-shaped inclusions of manganese sulfide, which formed during the production of the raw material.

The widths and depths of the melt tracks vary. The area of the base body 1.4301 shows strong line-shaped inclusions of manganese sulfide, which formed during the production of the raw material. The structure of the additive area is very different from that of a conventionally cast structure (Figure 3).AdditivelymanufacturedMaterial: 1.4404Base bodyMaterial: 1.4301Transition area50µmFigure 3: Micrograph of hybrid sample; macroscopic (left) and microscopic (right)In the connection area above the base body, material mixing is evident. This area is followed by a heat-affected zone in the base-body area, which shows a typical austenitic structure. The structure of the base body shows row-shaped delta-ferrite fractions. Compared to the heat-affected zone, the base body has a high proportion of deformation-induced martensite. Viewed from the edge, the heat-affected zone increases from 40 µm to 130 µm inside the component.

This behavior can also be seen in the test series, in which only one layer of powder is melted onto the base body (Figure 4).406100 µmBase body1.4301PowderfusionlayerPowderfusionlayer100 µmHeat–affectedzoneManganesesulfateBase body1.4301Figure 4: Polished (left) and etched (right) sample after exposure of one powder layerThe results of the microhardness tests show a hardness of 215 ± 3 HV0.05 in the additive range, which corresponds to a low load hardness value.Mixing AreaAdditive AreaHeat-Affected ZoneConventional Areamax.min.0501001502002503003504000246810121416Vickers hardnessHVMeasuringpointHardnessmeasurementofhybrid samples1.4404/1.4301 accordingtoVickers-HV0.05Figure 5: Vickers hardness measurement HV0.05 of the hybrid sampleIn the area of material mixing, the hardness values gradually increase and at the end of this zone they are within the hardness range of the 1.4301 material.

Hardness values of 320 ± 8 HV0.05 can be seen both in the heat-affected zone and in the conventional area (Figure 5).4074. DiscussionThe investigations using optical and scanning electron microscopes indicate that a complete connection of the two materials was achieved, since no cracks and only a very small degree of porosity are evident. This confirms the results of previous tests on tensile specimens [8]. The different widths and depths of the melt tracks are due to the different orientations of the exposure scan vectors, which differ for each layer.The difference in the microstructures of the additive and conventionally produced volumes is due to the layered structure. Due to the laser exposure it can be assumed that the microstructure is oriented in the direction of the highest temperature gradient. [14].

Due to the laser exposure it can be assumed that the microstructure is oriented in the direction of the highest temperature gradient. [14]. The layered exposures remelt the solidified material and, depending on the orientation of the scan vector, lead to an uneven orientation of the melt traces.The material mixing in the connection area can be explained by the dynamics of the melt pool. During the exposure of the first powder layer, the base body absorbs part of the laser radiation, melting it and releasing components of the 1.4301 material. Due to the dynamic of the melt pool, the melt of the build-up body mixes with those of the base body [15]. The optical microscopic investigations suggest that with each new exposure the alloy components of the underlying layer mix with the top layer and are carried upwards.

The optical microscopic investigations suggest that with each new exposure the alloy components of the underlying layer mix with the top layer and are carried upwards. The degree of this material mixing is reduced with each new layer of material applied.In the base body, microstructural changes in the connection area and heat-affected zone are evident. Since deposits of deformation-induced martensite form in the untreated state of the 1.4301 base material, the heat-affected zone can be identified by the area free of deformation-induced martensite, which is converted to austenite by the temperature effect of the build-up process.Compared to multi-layer exposure, the depth of the heat-affected zone is significantly lower with a single melted layer due to the lower heat input. The increased hardness in the connection area can be explained by the mixing of the two materials, wherein an inhomogeneous material distribution occurs [7].

The increased hardness in the connection area can be explained by the mixing of the two materials, wherein an inhomogeneous material distribution occurs [7]. Despite the changed microstructure, the lack of deformation- induced martensite does not affect the hardness of the heat-affected zone. It can be assumed that precipitation hardening leads to a compensation hardness in this area. Thus, laser processing results in changed hardness values only in the area of material mixing.5. Summary and outlookIn this work, the connection of hybrid components using different materials was investigated. 1.4404 steel material was used for the additively manufactured body and 1.4301 steel for the conventionally manufactured base body. The samples were examined by optical microscopy, scanning electron microscopy and hardness measurement. The following conclusions can be drawn from the results:A complete connection was made between hybrid components of comparable materials.

The following conclusions can be drawn from the results:A complete connection was made between hybrid components of comparable materials. No cracks or relevant pores were found.The connection area of the hybrid samples can be divided into the four zones: purely additive, material mixing, heat-affected, and conventional areas. The material mixing occurs in response to the dynamics of the melt pool.408During the design and execution of the tests it was determined that both materials must be compatible in order to ensure a crack and pore-free connection. Furthermore, the connecting surface on the base body must have low roughness and be parallel with the coating plane. In addition, exact positioning on the base plate and a strictly defined height of the first powder layer are required. Future research will focus on which material combinations prove most effective in the production of hybrid samples.

Future research will focus on which material combinations prove most effective in the production of hybrid samples. Notched bar impact tests should be conducted to enable greater understanding of the mechanical behavior of hybrid components.AcknowledgmentsThis work was funded by the Federal Ministry of Education and Research (funding code02P16K021). The authors are grateful for this support.References[1] Wohlers T (2018) Wohlers Report: 3d printing and additive manufacturing state of theindustry. Wohlers Associates, Fort Collins[2] ASTM F2792-12a (2012) Terminology for Additive Manufacturing Technologies, WestConshohocken, PA[3] VDMA - Arbeitsgemeinschaft Additive Manufacturing (2017) Technology Scout 2017:Technologien, Anwendung und Hersteller, Frankfurt[4] BARTOLOMEU F, Bucuimeanu M, PINTO E et al. (2017) 316L stainless steel mechanical and tribological behavior—A comparison between selective laser melting, hot pressing and conventional casting.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperExamination of the LPBF Process by Means of Thermal Imaging for the Development of a Geometric-specific Process Control T. Pichler*, J.H. Schleifenbaum*† *Fraunhofer Institute for Laser Technology ILT, Aachen (Germany) † Chair for Digital Additive Production DAP at RWTH Aachen University (Germany)AbstractThe development of process parameters for the Laser Powder Bed Fusion (LPBF) process is typically carried out by the manufacturing and metallurgical analysis of geometrically primitive test specimens (e.g. cubes). The process parameters identified in this way are used for the manufacturing of parts which are characterized by a high geometric complexity and a combination of solid and filigree component areas.

The process parameters identified in this way are used for the manufacturing of parts which are characterized by a high geometric complexity and a combination of solid and filigree component areas. Due to the discrepancy between the parameter development on primitive specimens and applications with complex parts, a geometric-specific process control is to be developed. In the context of this work different sample geometries are manufactured from Ti6Al4V by LPBF and the process is monitored by thermal imaging. The influence between component geometry and process parameters on the thermal behavior is shown.IntroductionIn the Laser Powder Bed Fusion (LPBF) process, parts are additively manufactured from metal powder. Powder layers with a thickness of typically 20 μm to 100 μm are selectively melted using laser radiation and locally solidified. The repetitive process of powder application and selective melting produces parts in layers.

The repetitive process of powder application and selective melting produces parts in layers. Due to the additive manufacturing principle, components of almost unlimited complexity can be manufactured. LPBF is used in numerous industries such as the automotive industry, aerospace, medical technology and turbo machinery. [1]The deficiencies of the LPBF process include the build time, which sometimes takes several days to complete, and the relatively low process robustness. [2, 3] In addition to an adaptation of the machine technology, there is the research question of whether these aspects can also be addressed by a further development of the actual LPBF process. One possible approach is to adapt the LPBF process control to the part geometry.In the following, an approach for adapting the LPBF process control to the pat geometry using the material Ti-6-Al4-V as an example is discussed.

Experiments on the influence of geometry in the area of a layer of the component (2D-Area) and in the area of already built component areas (3D-Area) are presented as examples. The investigations are carried out using a thermographic camera to display the temperature field within the LPBF process zone.State of the ArtApproaches for the adaptation of the process control to the component geometry were examined e.g. by means of the High Power SLM. The component is divided into a skin and a core area. The core area is processed with a higher layer thickness (typically 2-3 times higher) as well as a laser beam with TopHat intensity profile and a diameter of 500 μm to 1000 μm at laser powers in the kW range in order to achieve the highest possible build-up rate. In order to maintain the1511highest possible surface quality, the skin area is processed with a single-mode fiber laser with a maximum nominal power output of 400 W and a laser beam diameter of approx. 90 μm.

90 μm. [2]The system manufacturer EOS GmbH enables users of its EOSPrint 2.0 software to adjust the laser power as a function of the scan vector length using a scaling factor. [4]With this approach, overheating can be avoided, e.g. by reducing the laser power of short vectors in narrow areas of the part.Using the Advanced ToolPath Utility of the AutoDesk Netfabb software, a JavaScript programming interface can be used to individually control path creation. Depending on the machine type, the process parameters can also be adjusted [5]. Compared to the current state of the art, the software enables an extended adaptation of the process parameters and the build strategies, so that different areas of the part can be processed with adapted process strategies.In [6] a thermographic camera is used to monitor the LPBF process in order to automatically detect spatter particles in the process zone.

In addition, the effect of a variation of the process parameters laser power, scanning speed and hatch distance on the emission detected by the thermographic camera is displayed. The change in the length of the scan vectors induced by the scan strategy and the resulting temperature field with overheating effects when the scan vector length is shortened are also shown as examples.In [7], a thermographic camera under the processing of the material IN718 is used to determine the dimensions of the melt pool as a function of the scanning speed. A reduction of the melt pool width with an increase of the scanning speed is observed. A correlation between scanning speed and length of the melt pool cannot be proven.ApproachGeometric-specific Process ControlA prerequisite for the development of a geometric-specific process control for the LPBF process is its applicability to any components and geometry types.

The approach presented here is to adapt the LPBF parameters on two levels: On the one hand by analyzing the currently exposed layer, the so-called 2D-Area, and on the other hand by analyzing the already built component areas, the so-called 3D-Area. LPBF process parameters are then developed for both areas by manufacturing and analysing specific test specimens.1512Part to be printed2D Area3D Area(current layer)(already built region).. Ly..zLoose Powder (0°)Overhang (>0°)xxcross-section constrictionzyxFigure 1: Approach of the geometric-specific process control2D-AreaThe exposure of a layer within LPBF is characterized by vector paths that are traced with a typically focused laser beam at a given power and speed. Depending on the geometry of the component, this results in a profile of temporal and spacial energy input.

Depending on the geometry of the component, this results in a profile of temporal and spacial energy input. For the description of this temporal and spacial energy input, the repetition time can be used as a key figure: It describes the time that elapses until a point within the layer is reheated by its adjacent melt track. Figure 2 shows false color images of one layer in different components.barscubetrianglecomplex part;;-----?!i ~~1111111'111111'111 111111111111111J111111 111111111111,111111 111111,11111111--1111111111111111 I~llilll"""'IllTRep <= 4 msFigure 2: Repetition times of different components in false colors representationTTRep > 4 msfeed motion►The current layer is processed meanderingly by the laser beam in the direction of the indicated feed motion. Areas with a repetition time less than or equal to four milliseconds are marked red, areas with a repetition time greater than four milliseconds are marked yellow.

Areas with a repetition time less than or equal to four milliseconds are marked red, areas with a repetition time greater than four milliseconds are marked yellow. The repetition time is used to identify those areas within a layer where overheating is likely and adjustment of process parameters could be beneficial. In this example, the repetition time was initially chosen arbitrarily to explain how the model works.

In this example, the repetition time was initially chosen arbitrarily to explain how the model works. On an experimental level, the critical repetition time must be determined.3D-AreaFor the investigation of already built component areas, the component can first be dividedinto a so-called voxel (volumetric pixel) model, as shown in Figure 3.1513zCAD model of partyxvoxel model of partPossible properties of each voxelZ distance to powder - 2D distance to component contour - „Voxel density“ in specific regions of the partFigure 3: CAD and corresponding voxel model of a part to be manufactured by LPBFProperties can then be assigned to each voxel to characterize the geometric properties of the part at the corresponding location. Possible properties of a voxel can be e.g. the distance to the powder in Z-direction, the 2D- and 3D-distance to the component contour or the voxel density as a measure for the filigree of the component.

the distance to the powder in Z-direction, the 2D- and 3D-distance to the component contour or the voxel density as a measure for the filigree of the component. The LPBF process parameters can then be adjusted depending on the voxel properties. For example, for filigree areas with a relatively low voxel density, the energy input can be reduced by decreasing the laser power or increasing the scanning speed in order to avoid overheating.MethodologyIn the following, various samples for the 2D range and the 3D range will be manufacturedand the process emissions recorded with a thermal imaging camera.Experimental SetupThe experiments will be performed on an LPBF laboratory machine. Figure 4 shows thetechnical specifications and a picture of the system.Build AreaBuild platform diameter: 98 mm Build height: 150 mmPowder MaterialTi6Al4VLaserSinglemode Fibre Laser Max.

Power: 400 W (cw) Wavelength: 1080 nmParticle Size distribution15 μm – 45 μmOptical SystemScanlab HurryScan20, Scanlab VarioScan40 Spot Size: 75 μmPowder SupplierPraxAir Surface TechnologiesShielding gasArgon < 50 ppm O2 during processLaser Power* Scan Velocity* Hatch Distance*265 W 1200 mm/s 0.12 mmPowder depositionBottom-Loader, soft-recoater (silicone)Scanning Strategy*Meander, 90° increment, feed direction always against inert gas flowModificationchamber cover with viewing window for thermographic camera (ZnSe Window)Standard Parameters from preliminary experiments to achieve a porosity of less than 0.5 %Figure 4: Technical data and illustration of the LPBF laboratory machine, material data and standard process parameters15142D-AreaFor the 2D-Area 10 mm long webs of different widths (10 mm, 5 mm, 2.5 mm, 1 mm,0.5 mm) are manufactured (Figure 5).Scan Vectors (Laser on)········•Extended Vectors (Laser off)10 mm10 mmStandard5 mm1 mmyzyxx2.5 mm0.5 mmFeed Rate = Const.Figure 5: Schematic depiction of webs of different widths with variable and constant feed speedThe webs are manufactured with standard parameters on the one hand and with a constant feed rate on the other hand.

The constant feed rate corresponds to the feed rate for the web with the dimensions 10 mm X 10 mm. In order to implement the constant feed rate, the vectors of the shorter webs are extended, but the laser is switched off earlier, so that the desired web width is achieved. The extended vectors are dashed in Figure 5. The evaluation is carried out by qualitative examination and comparison of the false color representation of the thermographic image data.3D-AreaTo examine the 3D-Area, cube-shaped test bodies are placed on beams and the side length of the individual beams is varied to achieve different surface transition ratios (TR).

In the lower part a side view of the melt front of the web structures is shown, with the feed direction pointed into the image plane. On the basis of the web with a width of 10 mm, the melt front can be divided into 4 temperature zones: The laser zone, a subsequent heat zone and a subsequent cool down zone. Seen in the scanning direction in front of the laser-affected zone, the area of the powder that has not yet melted is also located.With the reduction of the web width, the characteristics of the melt front and the temperature zones change: First, the cooling zone becomes shorter until it is virtually non-existent at a web width of 2.5 mm. With this web width, only the laser influence zone, heat zone and powder zone are present. If the web width is further reduced to 1 mm and 0.5 mm respectively, the heat zone also disappears.

With this web width, only the laser influence zone, heat zone and powder zone are present. If the web width is further reduced to 1 mm and 0.5 mm respectively, the heat zone also disappears. Only the laser-affected zone and a very small area of the powder zone remain.With the reduction of the web width, in addition to the above described change in the characteristic of the melt front in the scanning direction or X-direction, a change in the characteristic in the feed direction or Y-direction also occurs. Figure 7 already shows that the edge1516steepness of the temperature field flattens in the Y-direction with a reduced web width.

Figure 7 already shows that the edge1516steepness of the temperature field flattens in the Y-direction with a reduced web width. Figure 8 shows the thermographic images of the web structures in top view.fr = 13.4 mm/sfr = 25 mm/sfr = 44.1 mm/sfr = 81.4 mm/sfr = 113.6 mm/sfr: feed rate [mm/s]; speed of movement of the laser spot in Y-direction (orthogonal to scan direction)Figure 8: Thermographic images of web structures in false color representation manufactured with standard parametersWith the reduction of the web width, the heat-affected zone is shifted further in the Y- direction: Whereas the heat-affected zone of the 10 mm wide web extends approx. 2.5 mm – 3 mm in the feed direction, the heat-affected zone of the relatively narrow 1 mm or 0.5 mm wide webs extends over the entire 10 mm Y-dimension of the web.

2.5 mm – 3 mm in the feed direction, the heat-affected zone of the relatively narrow 1 mm or 0.5 mm wide webs extends over the entire 10 mm Y-dimension of the web. The global movement direction of the laser spot shifts from the X-direction to the Y-direction with decreasing web width, as the increasing feed rate illustrates. The web with a width of 0.5 mm appears to have a similar characteristic to a single track, but with movement in the Y-direction.

The energy per unit length of the 0.5 mm web in the Y- direction is greater by a factor of 10 than the energy per unit length in the X-direction. This can be a possible explanation for the changed thermal profile as well as a potential source of overheating with correspondingly filigree structures.Figure 9 shows thermographic images of web structures, in the upper image area with standard parameters, in the lower image area with constant feed rate.

Areas with an emission value greater than 600 are colored red, areas with a lower emission value are colored grey.1517fr = 13.4 mm/sfr = 25 mm/sfr = 44.1 mm/sfr = 81.4 mm/sfr = 113.6 mm/sConstant feed rate for all webs: fr = 13.4 mm/sEmission2004006008001000Figure 9: Thermographic images of web structures, upper image area: standard parameters, lower image area: constant feed rate.In the upper row of Figure 9, the displacement of the heat-affected zone in the Y-direction with decreasing web width already visible in Figure 7 can be seen. In the lower row, the heat- affected zone is equalized by extending the scan vectors.

In the lower row, the heat- affected zone is equalized by extending the scan vectors. The expansion of the heat-affected zone in the Y-direction can be completely adjusted and is independent of the web width.3D-AreaFigure 10 shows the average emission values for the transition layer between the beam areaand the massive area of the specimen (layer 150) for different area transition ratios over time.1400Average emission of transition layer12001000TR = 0 % (Powder)TR = 16 %TR = 32 %n o i s s i m E800600TR = 64 %TR = 100% (Solid)400200005001000150020002500 Time [ms]30003500400045005000Figure 10: Average emission of transition layer for different transition ratios1518In addition to the cubes on beams, the courses of the emission values for the exposure of a solid cube and loose powder are also shown. All specimens rise to a peak value during the exposure in a timeframe of up to t = 860 ms, after which the cooling begins.

All specimens rise to a peak value during the exposure in a timeframe of up to t = 860 ms, after which the cooling begins. When comparing the two extreme cases, the peak emission value lies between 1240 (loose powder) and 390 (solid cube), and are thus more than a factor of 3 apart. The emission values of the cubes on bars lay in between of these extreme values, whereby the peak temperature drops with increasing surface transition ratio. As the surface transition ratio decreases, heating rate of the specimens increases during exposure. Especially for the cubes specimens with TR = 64 % and TR = 32 %, a kind of staircase effect is visible during exposure or heating. This results from the alternating exposure of areas with beams and areas with loose powder.

This results from the alternating exposure of areas with beams and areas with loose powder. It becomes clear that the heat dissipation in solid areas such as in the area of the beams or the solid cube is significantly greater than in areas with loose powder.Figure 11 shows the peak values of the mean emission over several layers, starting from the transition layer (layer 150) for the cubes on beams with different area transition ratios and for the solid cube as reference body. For metrological reasons, it was not possible to perform a measurement for each individual layer, so that the measurement points are separated by several layers. For all surface transition ratios it becomes clear that the relatively large overheating effects observed in the transition layer (layer 150) have already largely subsided after approx. 20 layers. With the layer thickness of 40 μm used in these experiments, this corresponds to a material thickness of 800 μm of solid material.

20 layers. With the layer thickness of 40 μm used in these experiments, this corresponds to a material thickness of 800 μm of solid material. In the further course of the layers, the emission peak values of almost all specimens approach those of the reference body. Only the test specimen with the lowest surface transition ratio of TR = 16 % shows a constant, albeit relatively small (less than 50 counts), increase in the emission peak value.1200Peak Values of average emissionTR = 16 %1000TR = 32 %TR = 64 %800TR = 100% (Solid)n o i s s i600m E4002000150170190210230250270 Layer290310330350370390transition layerFigure 11: Peak values of average emission in layers for different transition ratios1519In the experiment in Figure 10, a relatively strong overheating within the first 20 layers after the surface transition became particularly apparent.

The maximum emission value of initially approx. 1080 (laser power: 250 W) can be reduced to approx. 810 (laser power: 100 W). Nevertheless, the average peak emission value remains about twice as high as that of the solid reference body (approx. 400) in the previous experiments.15201521well as the identification of the relevant voxel properties for the implementation of geometric- Future work will aim at the determination of the critical repetition time in the 2D range asinvestigate the overheating effects in filigree component areas. constructing web structures of different widths. In the 3D area, cubes are produced on beams to On an experimental level, exemplary experiments in the 2D range are carried out byparameters according to the voxel properties. into account the different geometric properties of the component and to adjust the LPBF process the spacial and temporal energy input. The 3D area can be characterized by a voxel model to take 3D area.

The 3D area can be characterized by a voxel model to take 3D area. The 2D area can be represented by a repetition time model, which can be used to model method is presented. The components to be manufactured are to be divided into a 2D area and a An approach for the implementation of a geometric-specific process control for the LPBFConclusionemission. At this point a back-regulation to the reference power of 250 W seems to be expedient. (layer 170), almost all characteristic curves with reduced power undercut the reference range of the practice could involve the risk of porosity within the solidified component material. After 20 layers of the solid reference body. With increasing layer count, the standard range is undercut, which in After approx.

After 20 layers of the solid reference body. With increasing layer count, the standard range is undercut, which in After approx. 3 layers, the peak emission value at a power of 100 W is in the emission rangeFigure 13: Peak values of average emission in layers with variation of laser power for TR = 16 %170168166164162Layer 16015815615415215070200cube reference400600PL = 100 W~ tPL = 150 WPL = 200 WPL = 250 W8001000Variation ofLaser Power Peak Values ofaverageemission1200with TR = 16 %, varying the laser power for the first 20 layers after the transition layer.Figure 13 shows the course of the peak values of the average emission for the specimenE mi s s i o nspecific process control in the 3D range. Finally a geometric specific parameter map for the material Ti6Al4V shall be derived and demonstrated by the production of complex demonstrator components.Literature[1] Meiners, W.

amination of the powder spreading effect in Electron Beam Melting (EBM)C. Eschey, S. Lutzmann, M. F. Zaehiwb Institute for Machine Tools and Industrial Management, Technische Universitaet Muenchen, GermanyAbstractIn recent years, the scientific and industrial relevance of Electron Beam Melting (EBM) has grown. This is mainly due to the electron beam’s extensive power density and flexible positioning properties. Thus, considerable building rates as well as a favorable part quality can be realized. However, the appearance of transient physical effects constitutes a substantial drawback towards the broader use of the technology. Therefore, experimental examinations are being carried out in order to investigate the effect of sudden powder spreading during beam material interaction. Based on an existing mathematical model, an analytical approach is formulated in order to implement effective counter measures.

Based on an existing mathematical model, an analytical approach is formulated in order to implement effective counter measures. Hence, a significant increase in process stability is being achieved as the undesirable powder spreading effect is being avoided securely.IntroductionAdditive layer manufacturing is ideally suited to face current trends like increasing variety in versions at decreasing quantity. As the manufacture is done without a specific tool, the production of small batches can be conducted efficiently. The additive layer manufacturing for metal parts as well as for plastic parts is still dominated by systems using a laser beam source. Other technologies like electron beam melting still play only a negligible role. However, the use of an electron beam as an energy source for the selective melting of metal powder offers great potential like high beam spot velocity up to 5000 mm/s and beam power up to 10 kW.

However, the use of an electron beam as an energy source for the selective melting of metal powder offers great potential like high beam spot velocity up to 5000 mm/s and beam power up to 10 kW. Therefore, it is subject of current research projects.State of the artUp to now, only a few research efforts have been conducted concerning the process modeling of EBM. These few focus on the structured coverage of influence coefficients by [1] and the assumption of coarse approximate values for powder by [2]. The main topic of previous works is the explanation of an undesirable effect, which occurs strictly when the solidification step is being initiated. However, the mentioned research efforts do not handle this topic thoroughly. Through experiments with single layers of powder it was observed that under given circumstances the coated powder layer accelerates explosively inside the vacuum chamber [3].

Through experiments with single layers of powder it was observed that under given circumstances the coated powder layer accelerates explosively inside the vacuum chamber [3]. This so-called “powder spreading” effect appears mainly when powder particles are being exposed to the impact of an electron beam at room temperature. The effect exhibits a fundamental restriction on further developments of EBM [2]. With occurrence of this effect in the running process the last positioned powder layer will be destroyed and the layer has to berecoated. Figure 1 shows the powder spreading effect. The photographs were recorded by use of a high-speed-camera.t=0.000 st=0.040 st=0.160 st=0.240 s(a)(b)(c)(d)t=0.770 st=1.020 st=1.520 st=1.868 s(e)(f)(g)(h)Figure 1: High-speed recording of the powder spreading effectIn Figure 1(a) a powder layer on a base plate with a height of approx. 0.1 mm is shown.

0.1 mm is shown. The time t = 0.000 s denotes the moment at which the electron beam is being activated with a power of 100 W and first hits the powder layer. Figure 1(b) shows the powder layer at the time t = 0.040 s when the particles with a radius of 10 mm from the beam’s point of impact are in motion. Characteristic for that effect is that no large-scale melting of particles occurs [1] and that the beam is not modified in position or power respectively. After additional 0.120 s, Figure 1(c) shows the whole powder layer being accelerated especially in vertical direction. Therefore, only featureless gray boxes can be recognized in Figures 1(c) and 1(d). After t = 0.770 s (cf. Figure 1(e)), the base plate is no longer covered with particles. Instead, the powder is spread inside the vacuum chamber [3].Various mechanisms for the powder spreading effect have been examined in literature.

Instead, the powder is spread inside the vacuum chamber [3].Various mechanisms for the powder spreading effect have been examined in literature. With the posted approaches pulse transmission, thermodynamics, electrostatics and electrodynamics the aim of these explanatory models was to identify the dimension of active forces and so to identify the main contribution to the powder spreading effect. Based on that counteractive measures were designed in order to prevent the powder spreading [3, 4, 5].Experimental analysisDuring interaction of an electron beam with metal a significant fraction of the carried charge is assigned to the irradiated object, the so-called target. If the target is not electrically isolated, the charge is conducted completely to the grounding and no charging effect will take place [6]. According to this, a current through the ground wire can be measured, for example with a digital multimeter.

According to this, a current through the ground wire can be measured, for example with a digital multimeter. For a more detailed analysis of the powder spreading effect this current iMeas is acquired with a high-resolution measuring board (National Instruments PCI-4070) and synchronized with the recorded pictures of a high-speed-camera (RedLake MotionPro). For thisanalysis, the beam current IB is set to 1 mA and the position of the beam is not deflected. Under equal constraints two different properties for the irradiated surface are chosen. First, the electron beam hits the powder layer and second it hits the blank base plate. So it is possible to acquire the target current under the powder spreading effect as well as the current for the blank surface of compact material (cf. Figure 2). The solid line in Figure 2 shows the target current while hitting the blank base plate.

Figure 2). The solid line in Figure 2 shows the target current while hitting the blank base plate. As shown there is an almost constant target current of 0.69 mA, which decreases marginally after 1.1 s and shows higher fluctuation. Compared to that, the dashed line in Figure 2 shows the target current while hitting the powder layer. It is shown that this measured current has a distinctive volatility. After reaching the upper bound of approx. 0.67 mA, the graph declines to approx. 0.09 mA.

It is shown that this measured current has a distinctive volatility. After reaching the upper bound of approx. 0.67 mA, the graph declines to approx. 0.09 mA. Apart from small fluctuations, the current rises again after 0.4 s almost monotonically until the end of the measuring period after 2 s.Target current iMeas8start of powder accelerationx10-4 A654321000.40.81.2s2.0timereference measurement (target: solid steel plate)measurement during powder spreading effectFigure 2: Measurement of target current during powder spreading effect (blue dashed line)The high-speed-recordings during the powder spreading effect (cf. Figure 1) and the measured currents (cf. Figure 2) were taken concurrently. The frame rate was set to 1000 pictures per second. As it can be seen at time t = 0.240 s rather a powder cloud than a layer can be recognized, i. e. the intensity of the powder spreading effect is on its maximum degree.

As it can be seen at time t = 0.240 s rather a powder cloud than a layer can be recognized, i. e. the intensity of the powder spreading effect is on its maximum degree. At the same time the target current is minimal. Thus, it is obvious that the electron beam emits most of its charge by collision with the moving powder particles above the base plate. Only a reduced fraction of the inserted charge can be measured in the ground wire. The further pictures between t = 0.770 s and t = 1.868 s show, that the powder cloud decreases constantly in its intensity and thus, a greater amount of current flows into the base plate and into the ground wire. This monitoring correlates with the measured target current. The value of that current increases steadily up to iMeas > 0.6 mA.ModelingAn important observation through the above mentioned experiment is the 40 ms time-offset between the first contact of beam electrons with the powder surface and the start of powder acceleration (cf. Figure 2).

Figure 2). During this lag the powder particles are being charged negatively by the electron beam [6]. The particles’ amount of negative charge is crucial to the irradiated material’s electrical properties, e. g. electrical conductivity. Furthermore, the charge of the particles depends on the electrical isolation between the particles and the surface of the base plate [7]. However, it seems to be a proper description to split the powder spreading effect into a charge sequence and a discharge sequence (see Figure 3).electron beampowder particlesbase plateCharge sequenceDischarge sequenceFigure 3: Time-sequences of the powder spreading effectThe charge sequence lasts as long as the electron beam hits the powder and the single particles do not accelerate. With the upcoming spreading effect, the powder is no more under the pressure of the electron beam and thus only the discharge sequence is valid.

With the upcoming spreading effect, the powder is no more under the pressure of the electron beam and thus only the discharge sequence is valid. In the following paragraph, the procedure in the individual sequences will be investigated. An equivalent circuit diagram for the particularly relevant charge sequence will be introduced.Charge sequenceDue to extremely small contact areas between the single particles, metal powder generally has a considerably high electrical resistance [2]. This value is increased even more at the intersection between powder layer and base plate. This is mostly due to the base plate’s plain surface as there is only an exceptionally small contact area between the powder and the base plate [3]. This leads to the effect that the inserted charge cannot flow completely off through the grounded base plate and so it partially remains within the powder layer.

This leads to the effect that the inserted charge cannot flow completely off through the grounded base plate and so it partially remains within the powder layer. Because of this, the powder layer can be assumed as a parallel electrical connection between a resistor and a capacitor [8]. Figure 4 shows the idealized equivalent circuit diagram for the system electron beam/powder/base plate.IB,effRP,BPCP,BPuC(t)iC(t)iR(t)RBPRP,BP >> RBPAFigure 4: Equivalent circuit diagram for the system electron beam/powder/base plate during the charge sequenceThe amount of beam current is given by IB. Under consideration of the reflected fraction of charge carriers on the surface, the current IB,eff is inserted into the system. The above mentioned parallel connection of RP,BP and CP,BP is in series with the base plate’s resistor RBP. The beam current first has to flow through the powder and after that through the base plate to the ground wire.

The beam current first has to flow through the powder and after that through the base plate to the ground wire. The capacity CP,BP is considered as a parallel-plate capacitor. The powder serves as the negatively charged plate while the base plate serves as the positively charged plate. As mentioned above, the electrical resistance of the powder and its transition to the base plate RP,BP is significantly higher than the resistance of the base plate RBP [2].Due to the electrostatic interaction between a powder particle of the charge QP with other components of the system electron beam/powder/base plate it is exposed to different force influences by the law of electrostatic attraction [9]. First of all the force of repulsion has to be calculated, which a certain powder particle experiences through its interaction with a multitude of further equally charged particles in the powder layer.

The force of repulsion FP,P consists of many separate components as expressed in the following equation: 2 tQ )( P 2 r iPn1 1   0Ft )(PP ,4i1 In eq. (1), the dielectric constant is denoted by ε0 with a value of 8.854·10-12 C/Vm [9]. The distance between the considered powder particle and an arbitrary further particle i is denoted by ri. By the following eq. (2), the electrostatic force of repulsion between the electron beam and an also negatively charged powder particle i with the distance to the beam axis rB,i can be caluclated: (2)1   0 tQQ )( 2 r iB ,BPFt )(PB ,4In eq. (2), the variable QB can also be expressed by the following definition:Q BIBdt(1)(3)Regardless of the two above mentioned forces of repulsion also a force of attraction arises.

(2), the variable QB can also be expressed by the following definition:Q BIBdt(1)(3)Regardless of the two above mentioned forces of repulsion also a force of attraction arises. This force holds during the charge sequence between the base plate as the positively charged capacitor plate and the powder as the negatively charged capacitor plate. The value of this force of attraction FBP,P is being calculated as follows:FPBP ,t )(41   0 tQQ )( 2 r iBP ,CPThe charge of a capacitor’s oppositional electrodes has different signs and is of identical value [10]. Thus in eq. (4), the value of QC equates the negative charge of the powder as well as the positive charge of the base plate. The distance between an arbitrary particle i of the powder layer and the center of the base plate is denoted by rBP,i.Additional to the mentioned electrostatic interdependencies, a mechanical force interacts between the powder particles in the area of impact.

The force takes effect due to the pulse transmission of the moving beam electrons. According to [5], the force has the potential to accelerate powder particles which are hit by the beam and thus overcome the acting force of friction.The impelling force Fe,P can be estimated in consideration of the velocity of the beam electrons vB as follows [11]:F, Pe vm oBI B q cm o01   11 Uq A 2 cm  0 0   2I B qIn eq. (5) the electron mass is denoted by m0, the electron charge by q. The velocity of light is labeled as c0. The acting force is assumed to be equally distributed on a circular area of impact Aimp with diameter dimp where the beam interacts with the powder. The electron impelling pressure pe,P on the irradiated powder particles can be estimated as follows:pPe ,F APe ,imp4 cmI   B 2  dq  impo 01  11Uq  A 2  cm 0 0  2As denoted in eq.

(1), (2) and (4), the electrostatic forces of repulsion and attraction respectively influence an arbitrary powder particle during the charge sequence. Furthermore, there is an electron impelling force in the area of impact, see eq. (6). Hence, a superposition which leads to a maximum amount of force on the powder particles arises in the area of impact. From that position the layer starts to spread. This assumption is documented by the analysis of high-speed- recordings (cf. Figure 2). It is distinguishable that the powder particles begin to spread initially from the center of the beam in radial direction. Additionally, the particles are also accelerated in vertical direction. This indicates that the powder layer is not hit by the beam directly and the equivalent circuit diagram (cf.

Additionally, the particles are also accelerated in vertical direction. This indicates that the powder layer is not hit by the beam directly and the equivalent circuit diagram (cf. Figure 4) for the charge sequence is not valid any more.(4)(5)(6)Discharge sequenceThe discharge sequence during the powder spreading effect is distinguished by the fact that the powder on top of the base plate is moving and no more powder is resting on the base plate. The powder particles accelerate in vertical as well as in radial direction (see Figure 2). As already mentioned, the accelerated powder particles collide with the beam electrons and rest beyond the base plate.Thus, the electron beam no longer impacts the powder layer, but couples its energy or charge respectively into the grounded base plate. As it will be shown in the following paragraph, it is much more complicated to depict an equivalent circuit diagram for the discharge sequence than for the charge sequence.

As it will be shown in the following paragraph, it is much more complicated to depict an equivalent circuit diagram for the discharge sequence than for the charge sequence. Additionally, during the discharge sequence the powder spreading effect has already started and thus, adequate counter measures related to the charge sequence need to be considered. As common to a capacitors’ discharge sequence, the oppositely charged electrodes equalize. Due to the effect of spreading particles of the system electron beam/powder/base plate, the negatively charged electrode releases its charge. The amount of QC is decreasing mainly through this effect. The beam electrons which are inserted into the base plate in spite of their collision with the spread particle, cause a direct charge sink to ground due to RP,BP >> RBP. The beam electrons compensate the positively charged base plate, respectively. As a result, no more charging of the residual powder layer takes place during the discharge sequence.

The beam electrons compensate the positively charged base plate, respectively. As a result, no more charging of the residual powder layer takes place during the discharge sequence. The value of QP is thus completely determined by the charge sequence.Since a current can be measured during powder spreading, it indicates that the electron beam hits the base plate in spite of the emerging powder cloud (see Figure 2). Compulsory, this leads to a charge equalization of the base plate which had been positively charged during the charge sequence. In conjunction with the high electrical resistance RP,BP it follows that the powder particles spread before giving their negative charge completely into the base plate. From this it follows that the mentioned electrostatic forces are valid during the discharge sequence as well.

From this it follows that the mentioned electrostatic forces are valid during the discharge sequence as well. Since the powder particles are in motion, the electron beam cannot apply a force on the powder layer and so the electron beam pressure force can be neglected.During the discharge sequence another mechanism, namely the Lorentz force influences the powder spreading. The fundamental condition for that is that a negatively charged powder particle is accelerated to the velocity v by the above mentioned force. When crossing the magnetic field induced by the electron beam at flux density B a Lorentz force FL on the powder particle is being generated [9]. The force causes a vertical acceleration, as noted in eq. (7). QBvQF  PLP  v   20I B R    In eq. (7), the diameter of a current-carrying conductor of equal power to the electron beam is denoted by R.

(7), the diameter of a current-carrying conductor of equal power to the electron beam is denoted by R. The magnetic flux density is denoted by µ0 which is equal to 4·π·10-7 Vs/Am.(7)Conclusion from the modelThe quantitative prediction of process parameters is very complex. Since for proper determination the total number of powder particles and their respective properties, such as geometry, particle size distribution and electrical properties have to be considered within the calculation. Hence, a considerable computing effort would result. Therefore, the presented analytical model is used in order to predict the qualitative behavior of the powder spreading effect.As a result of the previously mentioned modeling, qualitative estimations for possible input coefficients are given (cf. eq. (1), (2) and (4)). An important aspect of previous sections is the time-discretization of the powder spreading effect during the charge as well as during the discharge sequence.

eq. (1), (2) and (4)). An important aspect of previous sections is the time-discretization of the powder spreading effect during the charge as well as during the discharge sequence. For a proper understanding, the distinction between causative and acting forces during the persistent powder spreading effect is important. The focus of consideration is the prevention of the powder spreading effect.

The focus of consideration is the prevention of the powder spreading effect. Since the charge sequence is terminated with the upcoming spreading effect, the following explanations focus on the acting forces and pressure components (see Table 1).Value DescriptionEq.Charge sequenceDischarge sequenceBasic parameterFP,PFB,PFBP,Ppe,PElectrostatic force (particle/particle) Electrostatic force (beam/particle) Electrostatic force (base plate/particle) Electron impelling pressure(1)(2)(4)(6)xxxxxxxRP,BP, CP,BP, IB, riRP,BP, CP,BP, IB, UA, rB,iRP,BP, CP,BP, IB, rBP,iIB, UA, dimpFLLorentz force(7)xTable 1: Analysis of the action of force and its causative basic parametersIn this context, “basic parameter” denotes the variables, which have direct or indirect impact on the mentioned equations of force and pressure components, respectively.

It is notable that the electric resistance RP,BP and capacitance CP,BP of the system powder/base plate as well as the beam current IB appear especially common. For the sake of clarity the basic parameter IB,eff is reduced to IB, since a change of IB under equal preconditions leads to a direct variation of IB,eff. Furthermore, the acceleration voltage UA and the beam spot diameter dimp are relevant. It is notable, that table 1 contains further basic parameters. First, the value ri concerns the distance between the powder particle under consideration and a further particle i. Second, the distance between a powder particle i to the beam axis is denoted by rB,i. Third, the distance between an arbitrary powder particle i to the center of the base plate is described by rBP,i. These values are required in order to estimate the acting forces on a representative powder particle in eq. (1), (2) and (4).

These values are required in order to estimate the acting forces on a representative powder particle in eq. (1), (2) and (4). Hence, they do not constitute a layer property and thus, can be neglected for further evaluations. Since the considerations for prevention of the powder spreading effect are restrictedto force and pressure effects during the charge sequence, the basic parameters of the Lorentz force are disregarded, too.Experimental validationIn order to achieve a matching between the mathematical-physical modeling of the powder spreading effect and real conditions, an experimental validation has to be arranged. According to table 1, the parameters “beam current” IB, “accelerating voltage” UA, “beam spot diameter” dimp, “shape of the powder particles” and the “preheating temperature” TP are relevant to the occurrence of the powder spreading effect. Some of the basic parameters (IB, UA, dimp) are directly reflected in these factors.

Some of the basic parameters (IB, UA, dimp) are directly reflected in these factors. In terms of RP,BP and CP,BP, these influences are being considered by the preheating temperature TP as well as by the particle shape. Both parameters have an impact on the resulting values of RP,BP and CP,BP [2]. The focus of experimental design is to evaluate impacts on the undesirable powder spreading effect. For this purpose, the chosen input parameters IB, UA, dimp, TP as well as the shape of particles are modified in two steps (cf. Figure 5). The layer thickness is set to 0.1 mm and the particle diameter fraction is 10 µm to 60 µm for both types of powder.50 µm10 µm50 µmspherical powder particlesspattered powder particlesFigure 5: Scanning electron microscope recordings of different powder particles shapesWhen all combinations are handled in terms of a full-factorial experimental design, a total number of 25 = 32 parameter settings result.

The focus of the scientific experiments is to hold the beam current on the specified value over a fixed period of ten seconds for every single test. However, when powder spreading occurs, the beam is deactivated immediately. In case the powder spreading effect has not taken place within the predefined period, the single test is completed and the powder is deemed not to be spread. In order to limit complexity of the experiments, the electron beam is neither altered in position nor altered in power during the defined period. For all experiments, the beam impacts the powder layer in its center. During preheating of the powder, the strongly defocused beam is moved across the layer in the shape of a circular area. The diameter of the circular area is 80 mm and the figure is defined by 1000 points. The overall area of the circle is being exposed to the beam 500 times within a second. Meanwhile, the beam current is linearly increased at 0.2 mA/s starting from 0 mA.

The overall area of the circle is being exposed to the beam 500 times within a second. Meanwhile, the beam current is linearly increased at 0.2 mA/s starting from 0 mA. This happens as long as the attached pyrometer indicates the adjusted preheating temperature TP of the powder layer. The used experimental design gives a review over the chosen settings (cf. Table 2). In order to eliminate unintentional impacts like trends, the order of the single experiments is defined randomly. Since the modification from one material to another involves a considerable effort, the overall experiment is divided into two parts. Within the first part (no. 1 to no. 16) the spattered10 µmpowder is examined, whereas in the second part (no. 17 to no. 32) the evaluations are limited to spherical powder. Therefore, the stochastic distribution of experiments is only valid within the particular sections. Besides the experimental design, table 2 shows the results of the single experiments.

Rather it is monitored that the powder starts spreading even during the preheating step. Hence, the experimental validation of the chosen beam properties is not possible. In contrast, the undesirable phenomenon never appears with spattered powder and a preheating temperature of 600 °C. As a further result, it is proven that an increasing defocusing of the electron beam leads to a positive effect. By increasing the beam spot diameter to a value of 0.4 mm, the powder spreading effect can be avoided in six single experiments. On the contrary, with a beam spot diameter of 0.2 mm this only holds true for four single experiments. Since the spreading effect can be avoided six times with IB = 0.5 mA and only four times with IB = 5 mA, the reduction of beam current is significant. However, theinfluence of the accelerating voltage is less relevant.

However, theinfluence of the accelerating voltage is less relevant. With both values of 60 kV and 100 kV respectively, spreading takes place three times each.ConclusionThe presented work relates to the increase of process stability for Electron Beam Melting (EBM). Up to date, the powder spreading phenomenon constitutes a considerable deficiency as the top layer of particles is being pushed away from the building plate. Hence, the deposition of another layer needs to be performed or the building job must be aborted respectively. Therefore, a modeling approach is being formulated concerning the examination of electrostatic, impelling electron and Lorentz forces. As a result, the respective influencing parameters were determined and thus, used for experimental validation with the EBM equipment.

As a result, the respective influencing parameters were determined and thus, used for experimental validation with the EBM equipment. It was detected, that the given equations sufficiently correlate with the conducted experiments.As pointed out in the previous section, the experimental results exhibit main influencing factors on the spreading effect. In this context, it is particularly notable, that the powder layer is able to be exposed to the electron beam at room temperature without the appearance of spreading when a beam spot diameter of 0.4 mm is being applied. Thus, the process step “preheating the powder layer” as it is state of the art in EBM can be neglected and a considerable saving of time can be achieved.

Thus, the process step “preheating the powder layer” as it is state of the art in EBM can be neglected and a considerable saving of time can be achieved. Furthermore, the efforts in terms of the removal of unmelted powder can be significantly reduced as the formation of sinter bonds is eliminated.The experiments also indicate that the powder spreading effect occurs permanently whenever spherical powder is being applied to the process. In reference to the previous modeling, the shape of the particles has a significant influence on the electrical resistance RP,BP as well as on the capacity CP,BP respectively. According to [3], spherical powder particles lead to an increase in electrical conductivity due to the fact that more and greater contact areas exist among the particles as well as between the particles and the base plate.

viewed PaperEXAMINING THE GPU ACCELERATION SPEED-UP FOR FINITE ELEMENT MODELING OF ADDITIVE MANUFACTURINGJ. Logan Betts1,2, Will Downs1,2, Matthew J. Dantin3 and Matthew W. Priddy1,21Department of Mechanical Engineering, Mississippi State University, MS, 39762 2Center for Advanced Vehicular Systems, Starkville, MS, 39759 3Naval Surface Warfare Center Carderock Division, West Bethesda, MD, 20817Keywords: GPU Acceleration, Additive Manufacturing, Finite Element Analysis, Directed Energy DepositionAbstractUsing a graphics processing unit (GPU) in addition to a central processing unit (CPU) has demonstrated promise for the acceleration of processing-intensive operations such as finite element (FE) simulations. Commercial FE solvers have begun to utilize GPU acceleration for classical multi-physics applications, but the speed-up for additive manufacturing (AM) simulations is not well understood.

Commercial FE solvers have begun to utilize GPU acceleration for classical multi-physics applications, but the speed-up for additive manufacturing (AM) simulations is not well understood. There is a significant need for GPU acceleration for metal-based AM FE simulations, which are computationally expensive because of the high mesh densities and the large number of time increments employed. This study examines the efficacy of GPU acceleration for Abaqus AM simulations, where benchmark simulations using a sequentially coupled FE thermo- mechanical model are run both with and without GPU acceleration. The speed-up is compared across the AM process for the thermal and mechanical analysis.

The speed-up is compared across the AM process for the thermal and mechanical analysis. In this study, GPU acceleration provided the ability to decrease simulation runtime by two-to-four times on 4-8 CPU cores, and one-to-two times on 16-32 CPU cores.IntroductionGraphical processing units (GPU) have been used in addition to a central processing unit (CPU) to accelerate highly parallelizable operations for over a decade [1]. The first paper detailing methods of using graphics hardware for matrix multiplication was published in 2001, but the first GPU designed for computing was not released until the release of the Nvidia G80 GPU in 2006 [2]. Alongside the G80, Nvidia released compute unified device architecture (CUDA), which allowed CUDA C/C++/FORTRAN language to be parallelized on GPUs [3]. The integration of CUDA allowed graphics hardware to be used in deep learning and supplied a path for GPU acceleration to multi-physics simulations.

The integration of CUDA allowed graphics hardware to be used in deep learning and supplied a path for GPU acceleration to multi-physics simulations. Limited support for GPU acceleration to commercial finite element analysis (FEA) solvers was introduced in Abaqus/Standard 6.11 in 2011, showing promise to decrease simulation runtime [4]. Early studies on using GPU acceleration focused on linear-static analysis, looking at decreasing the total of number of CPUs, thus requiring fewer acceleration tokens [4], [5].Metal-based additive manufacturing (AM) is a process for creating near-net shape parts layer-by- layer, using a moving heat source and material deposition. AM offers a reduction in processing and tooling steps when compared to traditional manufacturing, but often requires post-processing to reduce defects and residual stresses.

AM offers a reduction in processing and tooling steps when compared to traditional manufacturing, but often requires post-processing to reduce defects and residual stresses. Directed energy deposition (DED) is a metal-based AM process that uses a high-power coaxial laser with wire fed or blown powder deposition, but due to the localized heating and large thermal gradients experienced in the workpiece, the thermal history is hard to control and predict. Process parameters such as laser power, scanning speed, hatch spacing, and melt pool morphology drive the thermal history and affect the resultant microstructure and defects such as distortion, porosity, and residual stress formation.

FEA is commonly used to model and simulate AM processes to predict distortion and residual stress formation, as well as conducting parametric studies of printing parameters [6]–[9].Previous research on applying GPU acceleration to static mechanical analysis reported seeing up to a 2x speed up when using 1 GPU vs just 4 or 12 CPU cores [10], and a 5.2x speed up using 32 cores + 2x Nvidia K80 GPUs compared to 16 CPU cores using Abaqus. At present, no data could be found for using GPU acceleration with AM simulations in any commercial code, but custom codes have1561been used and showed speedups of 3-30x through a modified solution strategy or by relaxing constraints [11], [12].The purpose of this work is to present a first benchmark of the speedup of GPU acceleration applied to AM simulations in a commercial FEA solver. For the purposes of this paper, these simulations used Abaqus 2020 to model both an AM thermal and mechanical analysis of a 10-layer DED build.

For the purposes of this paper, these simulations used Abaqus 2020 to model both an AM thermal and mechanical analysis of a 10-layer DED build. As modeling AM processes with commercial codes become more prevalent for a cost-effective prediction of distortion and processing parameters, finding methods to decrease runtime of these computationally expensive simulations is critical.MethodologyThe AM thermal and mechanical analyses were run on a workstation equipped with an AMD Threadripper 3970x 32-core processor, 256GB of DDR4 Memory, and a Samsung 980 Pro 2TB NVMe storage device. The simulations were performed in three configurations: (i) no GPU, (ii) a Nvidia GTX Titan Black, and (iii) a Nvidia RTX 3090. The simulations were run on the workstation with 4, 8, 16, and 32 CPU cores for each case.

The simulations were run on the workstation with 4, 8, 16, and 32 CPU cores for each case. To compare the GPU acceleration performance to a traditional high performance computing cluster, the simulations were run on 20, 40, 80, and 200 CPU cores on a Cray CS300-LC cluster with two Intel Xeon Phi CPUs for a total of 20 cores per node.The thermal-mechanical analysis is conducted in two sequentially coupled analyses. First, the AM thermal simulation is performed to generate the temperature history of the part. Both the thermal and mechanical simulations model a ten-layer buildup of the DED process, specifically Laser Engineered Net Shaping (LENS) using Ti-6Al-4V as the material [7], [9], [13]. This specimen shown in Figure 1(a), is a 5 mm x 3 mm x 6 mm test specimen using a spiral scan strategy shown in Figure 1(b). The Goldak double ellipsoidal heat sourced is utilized, since a local moving heat source is the most computationally expensive method in the Abaqus/2020 [14].

The Goldak double ellipsoidal heat sourced is utilized, since a local moving heat source is the most computationally expensive method in the Abaqus/2020 [14]. Material deposition is accounted for using element activation [8]. The ten-layer benchmark utilized a seed size of 0.08 mm in the part corresponding to 147,537 DC3D8 (heat transfer analysis) and C3D8 (mechanical analysis) linear brick elements. Both analyses used a 0.025 s time step, 0.5 mm hatch spacing, and 16.6 mm/s scan speed. Temperature dependent properties including thermal conductivity, density, and specific heat are used. Additional information on the thermal analysis and thermal properties can be found in published literature [8], [9], [13].Figure 1: (a) dimensioned substrate and ten-layer benchmark specimen (b) laser scan strategy for each layerBefore the mechanical simulation starts, the temperature history is read in as shown in Figure 2.

The evolving microstructural model of inelasticity (EMMI) is the material model used to predict the mechanical response. EMMI is a flow rule based internal state variable (ISV) plasticity model [12]. An example of this implementation is shown as a flowchart in Figure 2. EMMI is ideal for DED due to1562temperature and rate dependence, cyclic temperature history, and recovery mechanisms that can relieve residual stress. The mechanical analysis can be used to make predictions of the residual stress or distortion within the part. More information on the calibration of material parameters and applications of EMMI for metal based AM is available in published literature [9].FEA simulations run in Abaqus require analysis tokens to perform an analysis. The number of Figure 2: An overview of the AM thermal-mechanical analysis framework using the commercial FEA solver Abaqustokens required depends on the number of CPU cores specified by the user and the user’s hardware.

To perform an analysis on a single CPU core, five analysis tokens are needed. An overview of the required number of analysis tokens with and without GPU acceleration is depicted in Table 1. Each GPU used behaves as a co-processor, only one additional analysis token is required to run a GPU alongside a CPU. The number of tokens required to run Abaqus on a number of CPU cores follows a logarithmic curve, and beyond eight CPU cores, no additional tokens are required to run a single GPU as a co-processor. The ability to implement GPU acceleration could allow users with few analysis tokens to run computationally expensive simulations efficiently.Table 1: Number of Abaqus tokens required for a given number of CPU cores. *For four CPU cores, one additional Abaqus token is needed for GPU acceleration.

*For four CPU cores, one additional Abaqus token is needed for GPU acceleration. Beyond eight CPU cores, one GPU can be included without additional Abaqus tokens.1563Number of CPUs 481620324080 200Abaqus Tokens8*12161721233146Results/DiscussionThe simulations were run with a range of CPU core counts, with the Nvidia RTX 3090 and GTX Titan GPUs, as well as without GPU acceleration. All other parameters were held constant, and no other simulations were run alongside any simulations, regardless of the number of CPUs used on the workstation. For the thermal analysis running the jobs with the GTX Titan yielded a speed up of 1.3x, 1.1x, 0.77x, and 0.68x for 4, 8, 16, and 32 CPUs, respectively compared to no-GPU accelerated runs. On the RTX 3090 the thermal analysis showed a speed up of 1.5x, 1.27x, 0.91x, and 0.82x for 4, 8, 16, and 32 CPUs, respectively, as depicted in Figure 3(a).

On the RTX 3090 the thermal analysis showed a speed up of 1.5x, 1.27x, 0.91x, and 0.82x for 4, 8, 16, and 32 CPUs, respectively, as depicted in Figure 3(a). For the mechanical analysis, the GTX Titan yielded a speed up of 2.98x, 1.8x, 1.8x, and 1.1x for 4, 8, 16, and 32 CPUs, respectively compared to no-GPU accelerated runs. On the RTX 3090 the mechanical analysis showed speed ups of 4.15x, 3.02x, 3.04x, and 1.82x for 4, 8, 16, and 32 CPUs as depicted below in Figure 3(b).The thermal and mechanical simulations were then run on a Cray CS300-LC cluster at 20, 40, 60, 80, and 200 CPUs without GPU acceleration. Similar runtimes of 2.75-3 hours were observed for the thermal simulation on 20-200 cores as depicted in Figure 3(c).

Similar runtimes of 2.75-3 hours were observed for the thermal simulation on 20-200 cores as depicted in Figure 3(c). Compared to running the thermalFigure 3: (a) AM thermal analysis comparison of runtime, (b) AM mechanical analysis comparison of runtime, (c) 4 & 32 CPU + 3090 GPU acceleration compared to cluster runs for thermal and (d) mechanical analysis.1564simulations on the workstation with 4 Cores with the RTX 3090, a slowdown of 0.96x was observed on average compared to the cluster runs for 20 to 200 CPUs. For 32 Cores with the RTX 3090, the thermal simulations saw a speed up of 1.31x, 1.44x, 1.33x, and 1.31x for 20, 40, 80, and 200 CPUs, respectively, on the Cray cluster. Similar to the speedup seen for no GPU to GPU, the mechanical analysis yielded the greatest speedup.

Similar to the speedup seen for no GPU to GPU, the mechanical analysis yielded the greatest speedup. For the mechanical analysis, the workstation with 32 CPUs and the RTX 3090 yielded a speed up of 1.7x, 1.3x, 1.1x, and 0.88x for 20, 40, 60, and 200 CPUs, respectively compared to cray cluster as depicted in Figure 3(d). The speedup for the mechanical simulations is summarized in Table 2.Table 2: Simulation speed-up comparing no-GPU to GPU Acceleration for mechanical analysisGPU32 Cores16 Cores8 Cores4 Cores30901.82x3.04x3.02x4.15xTitan1.1x1.8x1.8x2.98xThe nodal temperature for all nodes at each output time is extracted from the Abaqus output database and averaged for comparing the thermal response of the GPU and non-GPU accelerated simulations. An average percent difference of 0.99% was found between the GPU accelerated average nodal temperatures compared to the non-GPU accelerated simulations.

An average percent difference of 0.99% was found between the GPU accelerated average nodal temperatures compared to the non-GPU accelerated simulations. For the purposes of this initial study, this analysis serves as a verification, but further analysis is needed looking at how the parallelization of GPU acceleration will affect the results locally. It is understood that parallel computing will give slightly different results due to rounding errors when processing sums and products in different orders, even without GPU acceleration [15], [16].For the mechanical analysis, the workstation with 4 CPUs and the RTX 3090 yielded a speedup of 1.15x, 0.76x, 0.73x, and 0.6x for 20, 40, 80, and 200 CPUs, respectively compared to the Cray cluster as depicted in Figure 3(d). Implementing GPU acceleration allowed a simulation to be run faster on 4 cores using 9 acceleration tokens, than on 20 cores using 17 acceleration tokens.

Implementing GPU acceleration allowed a simulation to be run faster on 4 cores using 9 acceleration tokens, than on 20 cores using 17 acceleration tokens. Despite the simulations running 25% slower compared to 40 and 80 CPUs, the analysis used 14 and 22 fewer acceleration tokens. Without GPU acceleration, running AM simulations on 4 CPUs becomes impractical. Two implications of this work should be noted: GPU acceleration could enable AM simulations for users that (i) do not have access to high performance compute clusters or workstations and (ii) have limited access to acceleration tokens. Many universities or small research groups may not have access or funding to support compute clusters but could run these simulations on a small workstation with hardware like a gaming PC. While the reduction of 10 to 20 acceleration tokens may sound minimal, these acceleration tokens often cost thousands of dollars each, per year.

While the reduction of 10 to 20 acceleration tokens may sound minimal, these acceleration tokens often cost thousands of dollars each, per year. For smaller institutions, this difference of required tokens could be the difference in being able to run AM simulations at all.Between the thermal and mechanical simulations, the effect of GPU-acceleration varied. On 32- cores, the thermal simulation ran faster without GPU acceleration, while the mechanical simulation ran 1.82x faster on the RTX 3090 GPU. Our working hypothesis is that Abaqus only computes the average nodal temperature value (NT11) for 3D elements, which leads to fewer degrees of freedom (DOF) in the thermal analysis. This means that at each node in the thermal analysis, there is only one DOF, while in the mechanical analysis, there are six DOF at each node.

This means that at each node in the thermal analysis, there is only one DOF, while in the mechanical analysis, there are six DOF at each node. Our current working hypothesis is that an increased number of elements will affect GPU acceleration for a thermal analysis, but further investigation is needed.1565ConclusionsIn this work, a benchmark of using GPU acceleration for simulating AM thermal and mechanical simulations was presented. GPU acceleration was successfully applied to a 10-layer DED build in the commercial FEA solver Abaqus and demonstrated significant speedup in runtime—especially at lower CPU core counts. This benchmark only analyzed one mesh with 147,537 elements, so further study is needed to analyze how increasing the number of elements per CPU affects speedup of GPU acceleration. One possible avenue of exploration is investigating how CPU and GPU parallelization affects rounding errors and slightly changes the local results.

One possible avenue of exploration is investigating how CPU and GPU parallelization affects rounding errors and slightly changes the local results. The ability to decrease simulation runtime by two-to-four times by using GPU computing on 4-8 CPU cores is significant because of the accessibility of this computer hardware compared to specialty workstations and clusters. Many users may not have access to large numbers of costly acceleration tokens or dedicated compute clusters, but GPU acceleration could provide a way to run AM simulations efficiently with minimal acceleration tokens and consumer grade computer hardware. Future work will focus on understanding how the number of elements run on a CPU affects speedup and building large scale parts.AcknowledgementsResearch was sponsored by the Army Research Laboratory and was accomplished under Cooperative Agreement Number W911NF-20-2-0206.

AMINING VARIABILITY IN THE MECHANICAL PROPERTIES OF PARTS MANUFACTURED VIA POLYJET DIRECT 3D PRINTINGMichael W. Barclift and Christopher B. WilliamsDesign, Research, and Education for Additive Manufacturing Systems Laboratory Department of Mechanical Engineering, Virginia TechABSTRACTIn Objet’s PolyJet process, part layers are created by selectively inkjetting photopolymers onto a build substrate and then cured with ultraviolet lamps. With an eye towards using PolyJet as a manufacturing process to fabricate end-use products, the authors examine the sensitivity of part material properties to variation in process parameters. Specifically, a design of experiments is conducted using a full-factorial design to analyze the effects of three parameters on the specimens’ tensile strength and tensile modulus: the in-build plane part orientation (X-Y), the out-of-build plane part orientation (Z), and the distance between specimens.

Results show that part spacing has the largest effect on the tensile strength, but the three parameters produced no statistically significant effects on the tensile modulus. Orienting specimens in XZ orientation with minimal part spacing resulted in the highest tensile strength and modulus. Whereas, orienting specimens in the YZ orientation at the farthest part spacing led to the lowest mechanical properties.Keywords: Objet, PolyJet, Design of Experiments, 3D Printing1. INTRODUCTIONPolyJet Direct 3D Printing (PJD-3DP) is an Additive Manufacturing (AM) process that was developed by Objet Geometries, Ltd (Israel). In the process, layers of an acrylic- based photopolymer are selectively jetted onto a build-tray via inkjet printing (Figure 1). The jetted photopolymer droplets are immediately cured with ultraviolet lamps that are mounted onto the print carriage.

The jetted photopolymer droplets are immediately cured with ultraviolet lamps that are mounted onto the print carriage. The differentiating functionality of the PJD-3DP process (for the Connex machine to line) simultaneously jet multiple materials with different mechanical (and optical) properties. jetting "TangoBlack" (an elastomeric material) and For example, by simultaneously "VeroWhite" (a rigid polymer), a designer is able to manufacture products with functionally graded material properties. As such, PJD-3DP offers design freedom in specifying both the placement and the properties of materials on a point-by-point basis.isits abilityFigure 1. The PolyJet Direct 3DP Process [1]876While primarily used as a process for fabricating prototypes, the design freedoms offered by the process offer the opportunity to produce end-use products with unique, tailored geometries and materials.

One example is Objet’s marketing of FullCure 630 material, which is specifically formulated for the production of end-use custom-fit hearing aids. To be suitable for use as a manufacturing platform, the technology’s structure/process/property relationships must be well established. However, as it is a relatively new AM process, research on this technology is still emerging.1.1. AM Structure/Process/Property RelationshipsIdentifying structure/process/property relationships is critical for every manufacturing process; especially in AM where anisotropic properties can emerge due to material/energy patterning techniques, the build orientation of the part, and the interfaces between layers. Such research has been a key component of the advancement of all AM processes. For example, Agarwala and co-authors investigated how build strategies affect part porosity, and thus structural quality, in Freeform Filament Fabrication (FFF) [2].

For example, Agarwala and co-authors investigated how build strategies affect part porosity, and thus structural quality, in Freeform Filament Fabrication (FFF) [2]. Ahn and coauthors discovered that poor interlaying bonding in FFF contributed to significantly reduced tensile strength when parts were pulled perpendicular to the direction of layer construction [3]. Researchers have demonstrated that parts created by the Laser Sintering (LS) process are also anisotropic, with parts created as XY or XZ (part orientation and coordinate system used in ASTM F2921) were the strongest and parts oriented as ZX or ZY were the weakest [4, 5].Existing research of the PJD-3DP process has primarily focused in characterizing the resultant parts. Pilipovic and coauthors note that their measured values for tensile strength are lower than that which is reported by Objet [6].

Pilipovic and coauthors note that their measured values for tensile strength are lower than that which is reported by Objet [6]. Udroui and Mihail experimentally determined that the “glossy surface” finishing option provided a higher quality surface roughness than the “matte surface” option [7]. Existing structure/process/property research is limited to Brajlih and coauthors’ investigation of algorithms for scaling parts to improve process accuracy [8] and Kesy and Kotlinksi’s exploration of the effects of part orientation on material properties [9]. While hardness was unaffected by orientation, both tensile strength and elongation at break were found to be lowest for parts created with XZ orientation.Although structure/process/property relationships for PJD-3DP have not yet been fully the defined, stereolithography (SL) process, due to their common photopolymerization approach.

Paul Jacob’s extensive characterization of the SL process connects laser-based UV exposure of photopolymer resins to the cure depth of single vector scan to the mechanical properties of the finished part [10]. Specifically, elastic modulus and tensile strength increase (to a certain point) with increased exposure. This fundamental process characterization, which identifies the residual stresses caused by curing shrinkage [11, 12] and the “print-through” error caused by accumulated exposure across several layers, is linked to, and has driven the design of, SL laser scanning patterns [13]. SL process understanding has also provided a means of interpreting part anisotropy caused by varying layer thickness [14] and part orientation [15-17].

SL process understanding has also provided a means of interpreting part anisotropy caused by varying layer thickness [14] and part orientation [15-17]. In [17] it was found that tensile samples oriented as XZ or YZ were statistically stronger than those printed flat on build tray as XY or YX due to differences in inter-layer bonding.there perhapsis an opportunityto draw from established research877in1.2. Research ScopeAlthough PJD-3DP and SL are functionally linked by their use of photopolymerization [18], the manner in which the UV irradiation is patterned is quite different. The vector scan-based approach of SL provides the ability to precisely pattern UV energy across the resin bath. The Objet process, however, selectively patterns resin while indiscriminately distributing UV energy via lamps mounted on the printing-block.

The Objet process, however, selectively patterns resin while indiscriminately distributing UV energy via lamps mounted on the printing-block. Given that mechanical properties of photopolymer parts are so closely linked to the UV exposure, gaining an understanding of how PJD-3DP part properties relate to process parameters is paramount to ensuring its relevance and reliability as a manufacturing platform.In this paper, the authors explore the process/property relationships of PJD-3DP via a design of experiments. Specifically, a full-factorial design is used to analyze the effects of three parameters on the specimens’ tensile strength and tensile modulus: (i) the in-build plane part orientation (X-Y), (ii) the out-of-build plane part orientation (Z), (iii) and the distance between specimens.

This work is guided by the research question, “What controllable factors cause variability in the tensile strength and tensile modulus of VeroWhite parts manufactured through PolyJet Direct 3D Printing?” The PolyJet process is detailed in Section 2. The experimental design is outlined in Section 3 and the experimental methods are presented in Section 4. Results and discussion are presented in Sections 5 and 6, respectively.2. POLYJET DIRECT 3D PRINTINGAs briefly described in Section 1, the PJD-3DP process features direct inkjet printing of photopolymer resin and subsequent curing by UV lamps. The Connex machine features eight print heads that can simultaneously print three materials: four heads are dedicated for jetting support material; the other four heads jet two distinct build materials. A small roller in the printing block flattens jetted droplets to provide an even surface for the deposition of subsequent layers.

A small roller in the printing block flattens jetted droplets to provide an even surface for the deposition of subsequent layers. The process offers a relatively high resolution in the X-Y build plane (600 dpi printing resolution [19]) at small layer thicknesses (32 m, or 16 m in “High Quality” mode [19]). To provide scaffolding and stability to jetted droplets, printed parts are completely encased by the support material. The top-most surfaces are also typically coated in support material to provide an even surface finish (known as “matte finish”); however, the operator can choose to leave the top surface uncoated (known as “glossy finish”). Once the part has finished printing, the support material is removed from the part via a high-pressure water jet.During printing, the print-block moves along the X-axis (Figure 2a) and deposits material in two back-and-forth translations (which constitutes a full “pass”).

Jetting occurs only in the first forward translation; the remaining translations are solely to complete polymer curing. After completing a full pass, printing continues in the next printing “path” along the Y-axis of the build tray (Figure 2b). The extents of each printing path are defined by the width of the printheads. For example, the tensile specimens arranged on the build tray in Figure 2a each span three printing paths. The print-block offsets itself within each path between printing layers to fill in any voids that may have occurred due to clogged print nozzles.878(a)(b)Figure 2. Objet PJD-3DP TerminologyAs translation between printing paths is a process bottleneck, the process’s CAM interface (Objet Studio) automatically places parts in the build tray such that the longest dimensions are aligned along the X-axis, near the print origin. Additional “build rules” for orienting and placing parts within the build tray are presented in Table 1.

Additional “build rules” for orienting and placing parts within the build tray are presented in Table 1. These rules are derived as methods for minimizing total build time, minimizing support material consumption, and improving part surface quality. These rules are used for developing and selecting test parameters for experimental analysis, described in Section 3.Table 1.

These rules are used for developing and selecting test parameters for experimental analysis, described in Section 3.Table 1. Objet Build Rules [19] RuleDescriptionEffectRationaleX-Y-ZTall- Left Recess -UpFine- SurfacePlace longest dimension along X-axis  Place intermediate dimension along Y-axis  Place smallest dimension in Z-axis Place tallest part on the left of build trayOrient indented surface features facing-upOrient finely detailed features facing-upReduced overall build timeReduced overall build time Reduced total part cost Improved surface quality and lower costReduce number of printing paths  Reduce total number of layersReduce translations needed by print- block to create part Reduce consumption of support material  Improved resolution along Z-axis  Reduce support materialconsumption3. EXPERIMENTAL DESIGN AND METHODS3.1.

EXPERIMENTAL DESIGN AND METHODS3.1. Parameter SelectionGiven that mechanical properties of photopolymer parts are closely linked to UV exposure received (Section 1.1) and the manner in which the PJD-3DP process broadly patterns UV irradiation (Section 2), the authors hypothesize that a part’s mechanical properties might be affected by build orientation and/or the spacing between parts across printing paths. While orientation in the X-Y and X-Z planes have been shown to cause variability in mechanical properties, the samples in that existing study were all printed simultaneously on the same tray, and thus did not account to variations in exposure levels during the print [9].879To identify process variability, the authors chose to employ a design of experiments for the PJD-3DP process. To properly analyze and select pertinent parameters, [19] and [20] were consulted based on their "Six Sigma" and quality control methodologies.

To properly analyze and select pertinent parameters, [19] and [20] were consulted based on their "Six Sigma" and quality control methodologies. An Ishikawa Diagram (aka Cause-and-Effect or Fishbone Diagram) was created based on the authors’ knowledge of the PJD-3DP process (Section 2), and included potential factors that could produce deviation in the mechanical properties of VeroWhite material, as seen in Figure 3. Though this diagram does not guarantee nor reveal the true root cause, it was believed that through its graphical and organized structure, it would highlight the family where the root cause existed.Figure 3. Ishikawa Diagram for PJD-3DP ProcessThough many branches of the Ishikawa Diagram presented factors that could affect the material properties of the specimen, the “Machine” branch was further analyzed due to its uniqueness to PJD-3DP.

Three parameters, believed to be the most pertinent and fundamental to creating PolyJet parts, were chosen for analysis in this study.

A description for the selected parameters (X-Y Orientation, Z Orientation, and Part Spacing), and the rationale for their selection (i.e., hypothesized effect), is provided in Figure 4.880Parameter: X-Y OrientationXYDescription: The in-plane build orientation of the part such that its length is parallel or perpendicular with respect to the front of the build tray.YX Hypothesis: Orienting parts across print-head paths, may lead to lower mechanical properties due to banding from discretized jetting nozzles.Parameter: Z-OrientationYX or XY (flat)YZ or XZ (angled)Description: The build direction orientation of the part such that the width, not the length, lays flat or angled with respect to the X-Y plane.Hypothesis: An increase in the number of layers causes curing “print-through” and thus increased mechanical properties for parts with widths aligned in the Z plane (as noted for SL in [17]).Parameter: Part SpacingTight Spacing Description: The overall spacing between parts in X-Y plane of the build-tray.Far SpacingHypothesis: Smaller spacing may lead to increased mechanical properties due to potential UV over-cure effect.Figure 4.

Experimental Parameters3.2. Design of ExperimentsThe goal of the experiments is to identify the variability in the mechanical properties (specifically, the tensile strength and modulus) of parts created using the VeroWhite resin. Objet specifies that Fullcure 830 Verowhite has a tensile Strength of 49.8 MPa and tensile modulus of 2495 MPa [22]. Following theory presented in [23], a three parameter and two-level full factorial design of experiments (DOE) was selected to analyze the effects of the three selected parameters. The two-level full factorial design was selected because the parameters could be easily varied at two discrete levels and statistically analyzed using only eight total experiments, making the DOE easy to regulate and execute due to low complexity.Each parameter was varied at a "high" (+ 1) and "low" (-1) value and their effect on the tensile strength and modulus was measured. The full-factorial experiment is shown in Table 2.

The full-factorial experiment is shown in Table 2. Each experiment was given a codified name to quickly identify the experiment and the levels of881the parameters. For example, “YXT” corresponds to an experiment with specimens oriented with lengths along the y-axis, widths along z-axis, and tight part spacing (Figure 5).Table 2. Full Factorial ExperimentParameters LevelsExperiment Code NameYXT YXF YZT YZF XYT XYF XZT XZFX-Y Orientation YX (-1) XY (+1) -1 -1 -1 -1 1 1 1 1Z-Orientation YX or XY (-1) YZ or XZ (+1) -1 -1 1 1 -1 -1 1 1Part SpacingTight (-1)Far (+1)1 1 -1 1 -1 1 -1 1Figure 5. Build platform layout for “YXT” (specimens oriented in the y-axis, flat z-orientation, and tight part spacing) experiment.3.3. Experimental ProcedureASTM D638-10 was followed to measure specimens’ tensile strength and modulus. This standard has relevancy in engineering design and enables comparison across processes, including existing AM process analysis [17].

This standard has relevancy in engineering design and enables comparison across processes, including existing AM process analysis [17]. Figure 6 provides dimensional data for the selected "Type 1" tensile test specimen. Specimens’ thicknesses and widths were measured using a Marathon Electronic Digital Caliper (accuracy of 0.01 mm) to monitor any variable resolution and discrepancy. The specimens’ masses were measured on a digital scale (accuracy of 0.01g) were monitored to determine variation.Figure 6. ASTM D638 Type 1 specimen dimensions882Each experiment (Table 2) was randomly chosen and printed using an Objet Connex 350. Parts were printed in “Digital Material Mode” with the “Matte” surface finish. Random selection prevents experimental bias and propagation of a consistent error [23]. Eight total build trays were printed in this study.

Random selection prevents experimental bias and propagation of a consistent error [23]. Eight total build trays were printed in this study. Five specimens were printed per build tray, with each build tray corresponding to a specific experiment with uniform orientation among all printed specimen (no mixed orientations). Following the completed print job, specimens were cleaned using a high pressure waterjet. Upon drying, the specimens were conditioned at 72˚F in a closed opaque plastic container to prevent excess light exposure until ready for testing. Since [13] and [21] identify “time” as a possible degradation effect (due to oxygen levels and ambient light), the authors attempted to maintain a constant time between printing and testing for all experiments with a minimum conditioning period of 12 hrs.

The authors also attempted to control the humidity of the conditioning specimens by storing the plastic container in metal cabinet located away from potential sources of moisture.Tensile tests were conducted in the Virginia Tech Material's Testing Laboratory using an Instron 4468 Tensile Testing Machine in accordance to D638 standard. Using MTS Systems TestWorks for data acquisition with the Instron machine, tests were conducted at a stable environment of 72˚F and tested at a speed of testing of 50 mm/min (constant strain rate) until specimens ruptured. TestWorks was used to measure the applied force and strain of the specimen, which was then used to calculate the tensile strength and tensile modulus. The Tensile modulus was calculated using the Tangent Modulus method. Specimen elongation was measured with an axial clip-on extensometer (50% max strain). Test specimens were not tempered.4. RESULTS4.1.

A similar effect is also shown in that experiments fabricated with XY and YX orientations had widths consistently in the 13.3-13.4 mm range compared to the D638 specification of 13.0 mm. This suggests that Z- orientation can affect the accuracy of a part's thicknesses and widths.883The highest mean tensile strength was found in samples in the XZT experiment; the lowest occurred in the XZF experiment. The highest relative precision occurred in XZT with one standard deviation of 0.47 MPa (1.24 % its mean tensile strength) and the lowest occurred in YZF at 2.62 MPa (10.8 % of its mean). In comparison to Objet's quoted Verowhite tensile strength of 49.8 MPa [22], XZT had the lowest discrepancy at 12.0 MPa (24.1% of the quote) and XZF had the highest discrepancy at 26.9 MPa (54.0% of the quote).

Figure 7 graphically summarizes the tensile strength data with one standard deviation and shows that experiments with the tight spacing consistently show higher mean tensile strength values than at far spacing.Figure 7. Mean tensile strengthFigure 8. Mean tensile modulusThe highest tensile modulus occurred in the XZT experiment at a mean of 1874 MPa; the lowest occurred in YZF at a mean of 1176 MPa. The highest relative precision occurred in YXT with a sample standard deviation of 32 MPa (2.16 % of its mean tensile strength); the lowest occurred in YZF at 232 MPa (19.7 % of its mean). In comparison to Objet's quoted Verowhite Tensile Modulus of 2495 MPa [22], XZT again had the lowest discrepancy at 621 MPa (24.9% of the quote) and YZF had the highest discrepancy at 1319 MPa (52.9% of the quote). Figure 8 graphically summarizes the tensile modulus data.

Figure 8 graphically summarizes the tensile modulus data. It is observed that all experiments with XZ or YZ orientation and tight part spacing showed higher mean tensile modulus values regardless of the orientation. A similar effect is shown in that the lowest mean tensile modulus occurred with specimen widths aligned with Z-axis with far part spacing regardless of In-Build Plane orientation.4.2. Statistical AnalysisUsing the experimental data, a factorial analysis was performed to determine the effects of aparameter and its interactions on tensile strength and modulus (Tables 4 and 5).884Table 5.

These plots are a statistical technique for graphically estimating which main and interaction effects have statistical significance. Assuming the validity of the Central Limit Theorem [23], measured effects from a consistent process will appear to be from a normal distribution therefore, a potentially statistically significant effect would be an outlier.42 Main and Interaction Effects DataProbabilityNormal Probability PlotPart SpacingZ OrientationX-Y Orientation6815025050 Main Effects and Interaction DataProbabilityNormal Probability PlotX-Y OrientationZ OrientationPart Spacing300100200Figure 9. Tensile strength normal probability plot Figure 10. Tensile modulus normal probability plotFor the tensile strength, the factorial analysis revealed that the largest absolute effect was the part spacing, which could vary the response by 8.6 MPa.

Tensile modulus normal probability plotFor the tensile strength, the factorial analysis revealed that the largest absolute effect was the part spacing, which could vary the response by 8.6 MPa. Figure 9 shows that a majority of the effects and interactions appear to be from the normal distribution but, the Part Spacing may be the only statistically significant effect on the Tensile Strength. For the tensile modulus, the interaction between the Z-orientation and the part spacing produced the largest absolute effect at 310 MPa. The normal probability plot shows that all effects and interactions appeared to be from the normal distribution suggesting that none of them may have statistically significant effects on885the tensile modulus. While the part spacing was the largest absolute main effect for both responses, the Z-orientation was the smallest absolute main effect for both.4.3.

While the part spacing was the largest absolute main effect for both responses, the Z-orientation was the smallest absolute main effect for both.4.3. Analysis of Variance (ANOVA)Given that the Z-orientation parameter had the smallest absolute main effect, and appeared to be from the normal distribution for both the tensile strength and modulus (Figures 9 and 10), it was assumed to be statistically insignificant. This assumption permits the treatment of the performed three parameter and two level full-factorial design as a pseudo-replicated two parameter and two level full-factorial design (i.e., YXT to YZT and YZF to YXF were considered replicated experiments) [23]. By using this pseudo-replication, Analysis of Variance (ANOVA) could be performed on the two remaining effects (X-Y orientation and Part Spacing) and their interactions.

By using this pseudo-replication, Analysis of Variance (ANOVA) could be performed on the two remaining effects (X-Y orientation and Part Spacing) and their interactions. ANOVA is a statistical technique in which the sum of squares and error terms are compared against an F-ratio statistic along with p-values to determine the statistical significance of the parameters.Using SAS JMP, a statistical analysis software, experimental data was used in a "Fit Model" to define a two-level and two parameter full factorial design to perform an ANOVA. The DOE was modeled using a "Standard Least Squares" behavior. Each effect and interaction had its F- ratio calculated and compared against the critical F-ratio calculated at an alpha-value of 0.05. The ANOVA for Tensile Strength and Modulus is shown in Figure 11 and 12 below.Figure 11. Tensile Strength ANOVAFigure 12. Tensile Strength ANOVA8864.4.

The ANOVA for Tensile Strength and Modulus is shown in Figure 11 and 12 below.Figure 11. Tensile Strength ANOVAFigure 12. Tensile Strength ANOVA8864.4. Interpreting ANOVAThe ANOVA analysis of tensile strength provided an R-squared value of 0.88 and a RMSE of 2.45%, indicating an adequate correlation between the experimental data and the model suggesting the model is sufficient for prediction, due to minimal residuals and random noise. The factorial analysis in Figure 11's "Parameter Estimates" shows that the part spacing's was the largest absolute effect at 8.63 MPa, (which is approximately equal to the value from earlier analysis (Table 5). Its F-ratio of 24.8 is the only value that exceeds the critical F-ratio of 9.60 and its 0.0076 p-value is nearly an order of magnitude smaller than the 0.05 alpha value.

Its F-ratio of 24.8 is the only value that exceeds the critical F-ratio of 9.60 and its 0.0076 p-value is nearly an order of magnitude smaller than the 0.05 alpha value. Based on these two criteria, part spacing is the only statistically significant effect on the tensile strength.The tensile modulus ANOVA has an R-squared value of 0.43 but a RMSE of 232% which indicates that large residuals and a very large amount of noise exists in the data, suggesting that the model is not very adequate to predict the data. Figure 12's "Parameter Estimates" show that the part spacing had the largest absolute effect at 122 MPa, which is not approximately equal to the value found in earlier factorial analysis (Table 6). This discrepancy is due to the interaction effect between the part spacing and Z-orientation, which is confounded with the part spacing in this ANOVA.

This discrepancy is due to the interaction effect between the part spacing and Z-orientation, which is confounded with the part spacing in this ANOVA. While part spacing's F-ratio of 2.22 does exceed the critical F-ratio of 1.02, further analysis reveals that its 0.210 p-value is greater than the alpha value. Despite having a large F- ratio, part spacing and other effects are not statistically significant due to p-values greater than the alpha value, indicating a higher risk of Type I error.5. DISCUSSION5.1. Analysis of ResultsSynthesizing the experimental data presented in Section 4, the following results areidentified:Part dimensions aligned in the X-Y plane were fabricated 200-400 m larger than designed (Table 3). This suggests that Z-orientation can affect the accuracy of a part's thicknesses and widths. This result is not surprising given that the Z-resolution of the PJD-3DP process (32-60 mm layer thickness) is better than the X-Y printing resolution (600 dpi).

This result is not surprising given that the Z-resolution of the PJD-3DP process (32-60 mm layer thickness) is better than the X-Y printing resolution (600 dpi). This is result is unique to the PJD-3DP process as most AM processes have poor dimensional accuracy for features aligned along the Z-axis due to the discretized nature of the layer-by-layer approach.XY and YX parts did not show any statistically significant effects on material performance. The authors’ hypothesis that YX oriented parts would be weaker due to jetting from discretized nozzles was found to be incorrect. XY and XZ parts were, on average, stronger; however, the results were not statistically significant.Parts oriented in the Z-plane did not show any statistically significant improvements in material performance.

The authors’ hypothesis that parts with XZ and YZ orientation would be stronger due to an increase in the number of layers (and thus lead to increased curing due to “print-through”, as noted for SL in [17]) was incorrect. Specimens with widths along the Z-axis were, on average, stronger; however the results were not statistically significant.887Part spacing in the X-Y plane showed statistically significant effects on material performance. As hypothesized by the authors, parts printed closer together in the X-Y plane were stronger than parts printed further apart. This increase in material properties is hypothesized to be related to the manner in which the PJD-3DP indiscriminately patterns UV light during processing. Thus, when printing multiple parts that span multiple print paths, UV irradiation from the printing-block can over-cure parts in paths adjacent to the current printing path.5.2.

Thus, when printing multiple parts that span multiple print paths, UV irradiation from the printing-block can over-cure parts in paths adjacent to the current printing path.5.2. Sources of Experimental ErrorThe weakness of DOE is that there are no immediately replicated results to verify and validate the completed DOE. While ANOVA helps eliminate unnecessary variables, its success is based on the quality and precision of the collected data, in which replication would provide the best insurance for determining the consistency of the data. Full-factorial DOE's provide the highest quality and resolution in an experiment but considering each combination of parameter level can be time-consuming and lead to excessive use of resources. While a pseudo-replication can enable ANOVA for two parameters (Section 4.4), this assumption removes a parameter which provides less precision and insight as to how the original three parameters truly interacted together.

For example, while factorial analysis did not suggest that Part Spacing and Z- orientation had statistical significance, they had a practical significance in that their interaction was the largest absolute effect on the tensile modulus.In addition to error in statistical techniques, the lifetime of parts (i.e., time between print start and tensile test) was not consistent between experiments. This is important because time is suspected to be a potential cause of polymer aging and degradation [21]. In addition, differences of relative humidity between experiments and testing were not explicitly controlled; humidity can affect mechanical properties since polymers are hygroscopic and could have softened surfaces due to ambient moisture [24].

Finally, it is noted that a potential source of differentiation between the measured values with that specified by Objet could be due to their use of ASTM D638-03 and D638-04 [25], which might have different test specimens and conditions for testing and/or the authors’ use of “Digital Material” mode instead of “High Quality” mode during sample fabrication.6. CLOSUREThe purpose of this study was to analyze variability in the mechanical properties of parts created by the PolyJet Direct 3D Printing process due to changes in process parameters. The authors employed a design of experiments to identify changes in part tensile strength and tensile modulus due to changes in X-Y orientation, Z-orientation, and part spacing in the X-Y plane.

The authors employed a design of experiments to identify changes in part tensile strength and tensile modulus due to changes in X-Y orientation, Z-orientation, and part spacing in the X-Y plane. ANOVA revealed that part spacing had a statistically significant effect on mechanical properties; specifically, parts printed closer together were stronger than those printed farther apart.Future studies will further investigate the effect of part spacing on mechanical properties. The authors will explore the hypothesis that the PJD-3DP process is over-curing parts across multiple print paths. An AM process wherein part quality is affected by its layout on the build platform could have significant impact on a designer’s ability to predict part performance – in888effect, part properties would be inconsistent across various build platforms (due to part count and/or size).7.

ACKNOWLEDGEMENTSThe authors acknowledge Daniel Dressner (Virginia Tech ME, ’12) for his preliminary investigation in effects of build orientation on mechanical properties in the PJD-3DP process. The authors also acknowledge Mac McCord of Virginia Tech Material's Testing Laboratory for his assistance with tensile testing.8. REFERENCES[1] "Objet Polyjet Process." Objet Geometries Ltd. Www.Objet.com, n.d. Web.<http://www.objet.com/products/polyjet_technology/>.[2] Agarwala, M. K., V. R. Jamalabad, N. A. Langrana, A. Safari, P. J. Whalen and S. C.Danforth, 1996, "Structural Quality of Parts Processed by Fused Deposition," Rapid Prototyping Journal, Vol. 2, No. 4, p. 4-19.[3] Ahn, S-H, M. Montero, D. Odell, S. Roundy and P. K. Wright, 2002, “Anisotropic materialproperties of fused deposition modeling ABS,” Rapid Prototyping Journal, Vol. 8, No. 4, p. 248 – 257.[4] Gibson, I. and D.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceExpanding Material Property Space Maps with Functionally Graded materials for Large Scale Additive ManufacturingZeke Sudbury1,2, Chad Duty2,3, Vlastimil Kunc2,31 Cincinnati Incorporated, 2 University of Tennessee at Knoxville, 3 Oak Ridge National LaboratoriesAbstractBig Area Additive Manufacturing (BAAM) is a large scale extrusion-based print system that exceeds the throughput of conventional printers by five hundred times. In addition, BAAM uses pelletized feedstocks, which allows for site-specific definition of material composition and provides an unprecedented variety of material options. This study applies Ashby’s concept of a material property space map to a variety of materials suitable for printing on BAAM.

This study applies Ashby’s concept of a material property space map to a variety of materials suitable for printing on BAAM. Ashby maps plot the performance of various materials across multiple parameters (such as strength, density, stiffness, etc) allowing for direct comparison of non-dimensional performance criteria. This study uses Ashby maps to identify opportunities for the use of functionally graded materials on BAAM to achieve structural performance not yet available with conventional printers and homogeneous materials.IntroductionAdditive manufacturing (AM), commonly called 3-D printing, is the process of creating athree-dimensional part with consecutive two-dimensional layers. The part is designed withcomputer aided design (CAD) software, and then divided into multiple layers through a virtual“slicing” software [1]. AM encompasses a variety of technologies, including selective lasersintering, stereo-lithography, and polymer extrusion [2].

AM encompasses a variety of technologies, including selective lasersintering, stereo-lithography, and polymer extrusion [2]. One of the most common forms of 3-Dprinting is fused filament fabrication (FFF). In this process, filaments of thermoplastics, such asacrylonitrile butadiene styrene (ABS), are forced into a heating element, where it melts and flowsthrough a die. A gantry system moves the deposition system in a plane, and deposits the materialin a specified computer numerically controlled (CNC) path. This process is repeated for each 2-D layer. These FFF systems are typically restricted to flows rates of 70 grams per hour and buildvolumes around 0.125 cubic meters [3].Large format additive manufacturing, such as Cincinnati Incorporated’s Big AreaAdditive Manufacturing (BAAM), is an industrialized FFF system that creates parts 500 timesfaster than traditional FFF systems.

Love et al demonstrates thatadding carbon fibers to the feedstock increases strength, increases stiffness, and lowers thecoefficient of thermal expansion [4]. In addition, pelletized thermoplastics are approximatelyfive times cheaper than their continuous filament counterpart. The extruder on the BAAM is acustom, single screw extruder with six heating zones and one compressed air-fed hopper. Thebarrel is approximately 60 cm long, and has a 2.5 cm diameter screw designed for processingfiber filled polymers. The extruder has a maximum speed of 400 revolutions per minute (RPM)[5-7].With a variety of polymer and filler combinations, Ashby material property space mapsare needed to compare the performance of materials across multiple parameters. Duty [8] reportsmechanical properties for BAAM printed parts.

Duty [8] reportsmechanical properties for BAAM printed parts. The authors show that when a flatteningmechanism is used, the stiffness (Young’s modulus) of a printed material can be equal to orgreater than the reported values from injection molded reference. A BAAM specific materialproperty space map has been created by using the injection molded reference values reported bySabic [9], shown in Figure 1. Most of the materials group around the middle of the plot, leavinga large gap for potential improvement. Fleck et al identifies three ways to fill these holes inmaterial property space [10-11]. One method is to control the architecture of the material bycreating hybrids, such as composite reinforced materials, foams, and lattice structure. Using AMto create multifunctional structures with internal lattices are ongoing projects by other authors[12-14].

BAAM specific material property space map. The area circled in red identifies potentially accessible region with functionally graded materialsA functionally graded material (FGM) is an intentionally inhomogeneous hybrid, in which theconstituents are gradually varied so that material properties of the hybrid continuously changeover the geometry of the part to achieve a desired improvement in performance such as weight,mechanical properties, thermal properties, or cost [15]. In addition, varying the volume fractionof the constituents in a gradual manner reduces the likelihood of an interface failure [16]. Oneparticular application of FGM is a sandwich panel, in which the structure grades from a highstrength material at the face sheet to a low density material at the core. Traditional sandwichpanels tend to fail from poor bonding mechanisms at the sharp transition between materials, but aFGM sandwich panel reduces the likelihood of interface failure [17].

Traditional sandwichpanels tend to fail from poor bonding mechanisms at the sharp transition between materials, but aFGM sandwich panel reduces the likelihood of interface failure [17]. The majority of FGMstudies rely on simulations of mechanical or thermal behavior because manufacturing of FGMSis difficult. However, additive manufacturing allows design freedom over the geometry of a partand the pelletized feedstock of BAAM allows material to be easily blended and compositionallyaltered during a print. This study uses BAAM to manufacture functionally graded sandwichbeams. An analysis is performed to quantify the cost and weight of the beams.Experimental ProcedureFGM sandwich panels were manufactured using the most common BAAM materials. Twentypercent by weight carbon fiber-reinforced ABS (CF-ABS) was used as the face sheet material.461Unfilled acrylonitrile butadiene styrene (ABS) was chosen as the core material.

Twentypercent by weight carbon fiber-reinforced ABS (CF-ABS) was used as the face sheet material.461Unfilled acrylonitrile butadiene styrene (ABS) was chosen as the core material. Materialproperties for the two materials are shown in Table 1Table 1. Material properties and cost for BAAM grade polymersMaterialDensity (kg/m^3)Young's Modulus (GPa)Cost per Kg ($)20% CF-ABS114012.9$11Neat ABS10402.02$2 .2Front viewSide viewThickness (Z-axis) = 1.65”Length = 36”Width = 1.55”Figure 2. Geometry of beamThe beam geometry is shown in Figure 2. Since BAAM uses a layer thickness of 3.81mm, abeam with eleven layers (41.91 mm) was chosen. The FGM beam was created by first depositinga solid layer of CF-ABS to make the face sheet of the sandwich panel. Then, increasing amountsof ABS were mixed with CF-ABS and deposited in layers until a pure ABS core was achieved inthe middle (layer 6).

Then, increasing amountsof ABS were mixed with CF-ABS and deposited in layers until a pure ABS core was achieved inthe middle (layer 6). The process was then reversed to grade the beam composition from pureABS at the core layer to CF-ABS at the top face sheet. The gradient pattern was created with thepower law function𝑉 = (2 ∗ 𝑧 ℎ2 − ℎ1)𝐾where V represents the volume fraction of ABS in a given layer numer, z. A K value of 1 resultsin a linear relationship between the volume fraction and layer number. A lower K value equatesto a higher volume fraction of ABS throughout the part, as shown in Figure 3.462ABS Volume Fraction per Layerk=0.5k=10.7500.6000.4500.3000.150h / z0.0000.1500.3000.4500.6000.7500.0000.2000.4000.6000.8001.000Volume Fraction of ABSFigure 3. Volume percent of ABS in the FGM sandwich panel at each layer.

Volume percent of ABS in the FGM sandwich panel at each layer. Z/h represents distance from the neutral axisFigure 3 is a traditional way of reporting the volume of fraction of material in a FGM. Thevariable z/h represents the distance of each layer from the neutral axis. Another method ofrepresenting the material content in each layer is shown in Figure 4. In Figure 4, it is clear thatwhen K=0.5 (Power Law), there is more ABS in the transitions layers.463Volume percent of ABS per layerLinear K=1Power Law K=0.5100%90%S B A80%70%f o t n e c r e P e m u l o V60%50%40%30%20%10%0%123456 Layer Number7891011Figure 4. Graph representing the volume fraction in percent of ABS in each layer for both the Linear and Power Law variationsThe amount of material needed for each layer was calculated by converting the volume percentfrom Figure 4 into weight percent.

The material for each layer was pre-mixed in discrete batches(Figure 5). A CNC tool path of the beam from Figure 2 was created. The material for layer 1 waspoured into the BAAM hopper and the print was started. After each layer, the print was paused.Material remaining in the extruder was purged into a waste bin. The next layer’s material wasadded to the hopper, and was purged until steady state flow was achieved. The print wasresumed, and the process was repeated for all layers. Thermal imaging was used to verify that thetemperature of the previous layer remained above the glass transition temperature of ABS whenthe subsequent layer was deposited. Four beams were printed: One pure CF-ABS, one pure ABS,one Power Law FGM, and one Linear FGM.464Figure 5. Picture of the pre-mixed material for each layer. The white pellets are ABS, and the black pellets are CF- ABSResultsThe printed beams are shown in Figure 6.

Picture of the pre-mixed material for each layer. The white pellets are ABS, and the black pellets are CF- ABSResultsThe printed beams are shown in Figure 6. The black material is CF-ABS and the white material is ABS. The gradient is most visible in the Power Law variation because more ABS is present.a) Linear FGMb) Power Law FGMFigure 6. Pictures of printed sandwich panel beams with functionally graded coreA cost and weight analysis of the beams were performed by first converting the known volumepercent of ABS and CF-ABS of each layer into an actual print volume. Then, using the densityvalues, the weight of ABS and CF-ABS in the specific layers were calculated. The price per kgof the ABS and CF-ABS were used to find the total cost of the beam. Figure 7 plots thecumulative material cost of each beam.

The price per kgof the ABS and CF-ABS were used to find the total cost of the beam. Figure 7 plots thecumulative material cost of each beam. The pure CF-ABS and pure ABS beams have a linearrelationship between cost and layer because each layer uses the same weight of material, whichdirectly correlates to cost. The two functionally graded beams have a non-linear relationshipbecause increasing amounts of the cheaper material (ABS) are added. Both of the functionally465graded beams show a large price decrease compared to the pure CF-ABS beam. Figure 8 appliesthe same cumulative analysis to the weight of each beam. In this figure, the difference in weightbetween the four beam types is inconsequential.

Figure 8 appliesthe same cumulative analysis to the weight of each beam. In this figure, the difference in weightbetween the four beam types is inconsequential. Whereas the cost of ABS is 80% less than CF-ABS, the weight of ABS is only 9% less than CF-ABS.Cumulative Cost ComparisonLinear FGM CF-ABSPower Law FGM Neat ABS60.0050.00t s o C e v i t a l40.0030.00u m u C20.0010.000.00123456 Layer Number7891011Figure 7. Material cost as a function of layer for all four variationsLinear FGMPower LawCumulative Weight ComparisonCF-ABSNeat ABS6000.005000.00) g (t h g i e4000.00We v i t a l3000.00u m u C2000.001000.000.00123456 Layer Number7891011Figure 8. Beam weights as a function of layer count for all four variations466A comparison of weight and cost for each beam is shown in Table 2.

Beam weights as a function of layer count for all four variations466A comparison of weight and cost for each beam is shown in Table 2. Due to the similarity ofdensities in the chosen materials, the weight savings is small compared to the pure CF-ABSbeam. However, the large discrepancy in price creates a large savings in cost compared to theCF-ABS beam.Percent differenceBeam TypeWeight (kg)Weight (lb)CostWeightCostCF-ABS4.7610.48$ 52.42Linear FGM4.5710.07$ 32.944.0%37.2%Power Law FGM4.519.94$ 27.105.2%48.3%Neat ABS4.349.56$ 9.56ConclusionsBig Area Additive Manufacturing is a large format fused deposition modeling style printer thatuses pelletized feedstock to create parts faster than normal additive systems. In addition, thepelletized feedstock allows users to easily change materials between layers, which can be used tocreate a functionally graded material.

In addition, thepelletized feedstock allows users to easily change materials between layers, which can be used tocreate a functionally graded material. Sandwich panels with functionally graded cores haveshown to improve performance, but most publications are limited to simulations due to thedifficulty in manufacturing FGM panels. This study demonstrated the manufacture of FGMsandwich panels with BAAM. A cost and weight analysis revealed that using a gradient core canreduce cost by 48% and weight by 5%. Future work will test the mechanical performance of thebeams. Additional beams will also be printed and tested using a material with a larger weightdiscrepancy for the core.AcknowledgementsThe authors extend a special thanks to Local Motors in Knoxville, TN for allowing time on theirBAAM system to print the functionally graded beams.467References1. “3D printing of polymer matrix composites: A review and prospective.” Wang X.

PANDING RAPID PROTOTYPING FOR ELECTRONIC SYSTEMS INTEGRATION OF ARBITRARY FORMAmit Lopes*, Misael Navarrete**, Francisco Medina*, Jeremy Palmer***, Eric MacDonald**, Ryan Wicker* *University of Texas at El Paso, W.M. Keck Border Biomedical Manufacturing and Engineering Laboratory, El Paso, Texas 79968 **University of Texas at El Paso, Electrical and Computer Engineering, El Paso, Texas 79968 ***Sandia National Laboratories, Albuquerque, NM, 87185-0958Reviewed, accepted September 14, 2006AbstractAn innovative method for rapid prototyping (RP) of electronic circuits with components characteristic of typical electronics applications was demonstrated using an enhanced version of a previously developed hybrid stereolithography (SL) and direct write (DW) system, where an existing SL machine was integrated with a three-axis DW fluid dispensing system for combined arbitrary form electronic systems manufacturing.

This paper presents initial efforts at embedding functional electronic circuits using the hybrid SL/DW system. A simple temperature-sensitive circuit was selected, which oscillated an LED at a frequency proportional to the temperature sensed by the thermistor. The circuit was designed to incorporate all the required electronic components within a 2.5” x 2” x 0.5” SL part. Electrical interconnects between electronic components were deposited on the SL part with a DW system using silver conductive ink lines. Several inks were deposited, cured, and tested on a variety of SL resin substrates, and the E 1660 ink (Ercon Inc, Wareham, MA) was selected due to its measured lowest average resistivity on the SL substrates. The finished circuit was compared with Printed Circuit Board (PCB) technology for functionality. The electronic components used here include a low voltage battery, LM 555 timer chip, resistors, a thermistor, capacitors, and Light Emitting Diodes (LEDs).

The electronic components used here include a low voltage battery, LM 555 timer chip, resistors, a thermistor, capacitors, and Light Emitting Diodes (LEDs). This circuit was selected because it (1) represented a simple circuit combining many typically used electronic components and thus provided a useful demonstration for integrated electronic systems manufacturing applicable to a wide variety of devices, and (2) provided an indication of the parasitic resistances and capacitances introduced by the fabrication process due to its sensitivity to manufacturing variation. The hybrid technology can help achieve significant size reductions, enable systems integration in atypical forms, a natural resistance to reverse engineering and possibly increase maximum operating temperatures of electronic circuits as compared to the traditional PCB process.

This research demonstrates the ability of the hybrid SL/DW technology for fabricating combined electronic systems for unique electronics applications in which arbitrary form is a requirement and traditional PCB technology cannot be used.Keywords: rapid prototyping; stereolithography; direct-write; hybrid manufacturing; functional integrated layered manufacturingIntroductionWith recent advances in Rapid Prototyping of High Density Circuitry (RPHDC) research and specifically the development of a hybrid stereolithography (SL) and direct-write (DW) system, three-dimensional (3D) structures can be fabricated with high-resolution electrically-644conductive media (Wicker et al., 2005). Consequently, electronic components can be tightly integrated together in a rugged, light-weight structure of arbitrary form.

Consequently, electronic components can be tightly integrated together in a rugged, light-weight structure of arbitrary form. To date, the capabilities of this new RP technology have not been explored in the context of fabricating and prototyping complex electronics applications in which conformal shape, rugged and light-weight integration and a natural resistance to reverse engineering are all of paramount importance. Examples of applications currently being explored include:(cid:120) (cid:120) (cid:120) (cid:120) (cid:120)unattended / wireless sensors combined electrical and structural components in unmanned aerial vehicles soldier-borne electronics pavement instrumentation implantable bio-electronicsThis paper describes initial efforts at evaluating the effectiveness of the hybrid SL/DW system for embedding a simple parametrically-sensitive system into a SL matrix.

The natural progression of this work includes the fabrication of electronics with more sophisticated semiconductor devices (i.e. microcontrollers, field programmable gate arrays, memories, sensors, etc.) for a wide range of applications.Solid freeform fabrication (SFF), layered manufacturing (LM), and other similar rapid prototyping (RP) technologies enable the fabrication of complex three-dimensional (3D) parts. Stereolithography (SL) is one of the most widely used RP technologies to manufacture highly complex and accurate 3D prototypes due to its high build resolution. The materials development for SL and other RP technologies, focused on improving the mechanical strength, durability, and thermal property characteristics of RP materials, has generated great interest in many areas of research such as rapid manufacturing of fully functional embedded electronic devices.

The flexibility of placing components within complex 3D designs and the possibility of layering components instead of placing them side-by-side can lead to a reduction in size of the overall circuit. The overall strength of the circuit can be increased and components are protected from outside hazards as compared to its equivalent PCB design. These were some of the unique features and motivating factors that led to this research.Researchers at the UTEP W.M. Keck Border Biomedical Manufacturing and Engineering Laboratory and Sandia National Laboratories have developed a hybrid manufacturing setup to assist in expanding the markets for RP into rapid manufacturing of such unique functional devices. In this integrated manufacturing environment, SL is employed in conjunction with direct write (DW) fluid dispensing technology to capitalize on each of the individual process capabilities for developing unique electromechanical devices.

In recent years, researchers have developed systems for automatic dispensing of conductive, thermally curable media, such as DW inks for maskless patterning of electronics. Palmer et al. (2004) proposed a new manufacturing process to integrate SL and DW technologies for the fabrication of 3D, high-density circuitry. Medina et al. (2005) utilized this hybrid SL/DW system to demonstrate an SL part with embedded circuitry and 3D connectivity. The electronics industry, where huge performance improvements in each generation of new products are accompanied by the need to continuously reduce costs, is under continuous pressure from consumers to come up with product and design innovations. A number of these innovations need to be applied to satisfy the increasing pressure on the interconnect industry (Mosses and Brackenridge, 2003).

A number of these innovations need to be applied to satisfy the increasing pressure on the interconnect industry (Mosses and Brackenridge, 2003). The most commonly used interconnects in electronic circuitry is the PCB, where high volume manufacturing can be expensive and wasteful in spite of being suitable for many applications. Embedding electrical components using RP technology provides fewer process steps enabling reduced manufacturing645costs (Begaye et al., 2003). RP also increases design flexibility and almost completely eliminates waste due to its additive process. Passive components such as inductors and capacitors are key technology for highly integrated electronic circuits, but typically require two- thirds of the space of a conventional PCB (Waffenschmidt et al., 2005). Investigating the feasibility of embedded sensors or functional components in a layered structure will reduce surface area which could lead to total assembly cost reductions.

Investigating the feasibility of embedded sensors or functional components in a layered structure will reduce surface area which could lead to total assembly cost reductions. RP electrical interconnect prototypes could find uses in pre-production testing and three-dimensional circuitry can lead to full departure from conventional circuit PCB boards (Ocampo et al., 2003).The current effort represents the continuing evolution of the Functional Integrated Layered Manufacturing (FILM) concept introduced in the work of Medina et al. (2005) where RP is employed in conjunction with automated access to a toolbox of additional manufacturing technologies to capitalize on individual process capabilities for developing unique functional devices. As the next step in this development, this work presents a hybrid SL/DW machine setup for manufacturing embedded electronic components in a semi-automated environment with DW interconnects.

As the next step in this development, this work presents a hybrid SL/DW machine setup for manufacturing embedded electronic components in a semi-automated environment with DW interconnects. The hybrid system provides a method for automated deposition of fluid media during the SL build (by stopping the SL build but not removing the part from the SL machine). The DW system employed a three-axis translation mechanism, using precision linear positioning stages, that was integrated with the commercial SL machine. The current system allows for easy exchange between SL and DW in order to manufacture fully functional 3D electrical circuits and structures in a semi-automated environment. To demonstrate the manufacturing capabilities, the hybrid setup was used to make a simple multi-layer SL part with embedded electronic components with DW interconnects.

To demonstrate the manufacturing capabilities, the hybrid setup was used to make a simple multi-layer SL part with embedded electronic components with DW interconnects. The following describes the design and demonstration of the semi-automated hybrid SL/DW machine for manufacturing embedded electronic components.Hybrid stereolithography/direct-write machine setupAs depicted in Figure 1, an integrated RP system comprising a traditional SL apparatus and a DW fluid dispensing system were used in the hybrid SL/DW manufacturing setup. Many of the details described in this section are also found in Medina et al. (2005). However, this section provides several upgrades to the previously described system and represents the latest developments for the hybrid SL/DW manufacturing system. The SL 250/50 machine manufactured by 3D systems was used with a 355 nm solid-state laser upgrade (Series 3500- SMPS, DPSS Lasers Inc., Santa Clara, CA).

The SL 250/50 machine manufactured by 3D systems was used with a 355 nm solid-state laser upgrade (Series 3500- SMPS, DPSS Lasers Inc., Santa Clara, CA). Due to its ability to withstand heat deflection temperatures (HDTs) up to 126 ˚C and relatively low viscosity (550 centipoise, cp, at 30 ˚C) (DSM Somos®, New Castle, Delaware) DSM Somos® ProtoThermTM 12120 was selected as the resin for this work. The low material viscosity of ProtoThermTM facilitated improved handling, building without sweeping, and cleaning prior to DW deposits. This was one of the issues encountered in the work of Palmer et al. (2004) where it was challenging to handle and build parts out of ProtoToolTM due to the high value of viscosity (2500 cps).The DW setup was placed on a custom-made, adjustable height table, which consisted of a stainless steel top and an adjustable-height hand-crank two legged mobile workbench (Part # 9054T113 and Part # 3517T82, respectively, McMaster-Carr Supply Company, Chicago, IL).

Two support plates were welded to the stainless steel table top which was then bolted to the workbench base using standard fasteners. The DW setup was aligned to the SLA 250/50 using two identical dowel pins and the support plate on the SLA 250/50 frame until both pins touched the rim thus arranging the X stage parallel to the rim of the SL machine.646The 3-axis movement was achieved by precision linear positioning stages with direct drive precision ground screw drives and built-in 0.1 μm resolution linear encoders. The X-axis motion was provided by a 400mm travel LM150D-0400 model, while the Y and Z axis motion was provided by the LM100D-0100 model (Parker Hannifin Corporation, Electromechanical Automation Division, Rohnert Park, California) with 350mm and 100mm travel, respectively. The X axis was attached to a standard Newport honeycomb breadboard (Newport Corporation, Mountain View, CA) using an aligning plate provided by Parker Hannifin Corporation.

The X axis was attached to a standard Newport honeycomb breadboard (Newport Corporation, Mountain View, CA) using an aligning plate provided by Parker Hannifin Corporation. The X, Y and Z axes were aligned perpendicular to each other using support brackets as shown in Figure 2b. The 3-axis traverse stages of the DW system facilitated movement of the fluid dispensing system, enabling precise fluid placement without intermediate calibration or removal of the SL part.Precision 3-axis translation stageDPSS solid state laser systemDispensing system console3D Systems 250/50 SL machineAdjustable height mobile workbenchTranslation stage controllerDW dispensing syringe and nozzleFigure 1. The hybrid SL/DW machine setup.The 2405 Ultra Dispensing System (EFD Inc., East Providence, Rhode Island) shown in Figure 2a was integrated with the SL machine and adapted to deposit fluid such as conductive ink on the part manufactured by the SL machine.

The dispensing system as described previously obtains an externally supplied inlet air pressure (up to 100 psi) and regulates the outlet pressure (from 0 – 5 psi), which is then supplied to the dispensing tip (Medina et al., 2005). The ink can be dispensed manually or controlled using LabVIEW® to draw given circuit profiles and adapted to dispense a variety of fluids.For this work, a conductive ink (E1660-136, Ercon, Wareham, MA) was dispensed on laser cured ProtoThermTM 12120 resin (DSM Somos®, New Castle, Delaware). The ink used here was selected based on its low curing temperature (138 ˚C for 15 minutes), and lowest measured647resistivity (0.011 (cid:525)/square/mil) based on ink dispensing experiments. The traverses as well as the dispensing system were controlled using LabVIEW®.Dispensing System consoleYZDispensing syringe with nozzleX(a) (b)Figure 2.

The traverses as well as the dispensing system were controlled using LabVIEW®.Dispensing System consoleYZDispensing syringe with nozzleX(a) (b)Figure 2. (a) The DW Ultra Dispensing System; b) The 3-axis traverse system.Figure 3 summarizes the results of the ink dispensing experiments. The substrates used for this experiment consisted of a standard FR-4 board along with SL samples, which were simple 3” X 3” X 0.25” parts fabricated on the SL 250/50 machine using the ProtoThermTM 12120, NanoFormTM 15120, and ProtoToolTM 20L (DSM Somos®, New Castle, Delaware) resins. The SL parts were cleaned using Isopropyl Alcohol and post-cured in a UV oven prior to the DW process. A set of five one inch DW lines using a variety of inks were deposited on these substrates and cured at the recommended curing conditions.

A set of five one inch DW lines using a variety of inks were deposited on these substrates and cured at the recommended curing conditions. A dispensing pressure of 4 psi and Z gap of 4 mils between the dispensing tip and the SL substrate were chosen for this experiment based on previous research and gave the most consistent DW line (Medina et al., 2005).0.060Recommended resistivity) l iM q S / s m h O/( y t i v i t s i s e R0.0500.0400.0300.0200.010Prototherm substrateFR4 substrateNanoform substratePrototool substrateSelected conductive ink0.000CI1002CI1001E1660 InksE1400PT10AFigure 3. Comparison of average resistivity of various inks on different substrates.The results indicate that the E 1660 silver conductive ink gave the lowest average resistivity after thermal cure on all the substrates as compared to the other inks and was thus selected for creating the electrical interconnects between the electronic components.

Future research will include experimentation with smaller dispensing tips at various dispensing speeds to get the narrowest consistent DW interconnect line at the fastest speed. The newly acquired precision linear positioning stages will enable this experimentation.648Hybrid SL/DW machine operation overview and L555 timer circuit design reviewFigure 4 provides an operation overview of the LM 555 timer circuits with embedded electrical components and DW interconnection between the components. This figure is intended to provide a general description of the fabrication process using the hybrid SL/DW system for embedding electronic components within SL parts.

This figure is intended to provide a general description of the fabrication process using the hybrid SL/DW system for embedding electronic components within SL parts. The figure represents separate STL files for inserting as well as embedding the electronic components using different SL parts, and DW electrical traces for connection between the electronic components.Sockets for embedding electronic componentsEmbedded components with access holesDW traces for electrical interconnectsDW start point(a) (b) (c) Figure 4. (a) SL part with sockets for embedding electronic components, (b) Embedded components with access holes for DW, (c) DW traces for electrical interconnects.The ability to accurately register the DW trace with the SL part and providing precise electrical interconnections using the DW system are some of the key elements of successful SL/DW integration.

The registration issue is solved by locating the dispensing tip to a predetermined reference point, which is a simple cylindrical hole on the SL part, as shown in Figure 4c. A vertical gap of 4 mils is maintained between the DW ink dispensing nozzle and the SL part to achieve a consistent DW line based on previous research (Medina et al., 2005). The SL vat is lowered to allow access to the sockets and vias for inserting the electronic components and during the DW process while the platform remains at the same z-height. This procedure eliminates the geometric differences between the elevator positions before and after the DW process. These features summarize the necessary elements for successful hybrid SL/DW operation. Additional details of the operation are provided below in the description of part manufacture illustrated in Figure 4.Figure 4 represents a simple part with embedded electronic components to be manufactured using the hybrid SL/DW setup.

Manufacturing issues like part re-registration, continuous cleaning, and curing between part builds, as mentioned in Medina et al. (2005) are addressed using the hybrid setup. SolidWorks modeling software or similar CAD packages can be used to generate the individual part designs, and converted into STL file format as required. 3D Systems proprietary slicing software, 3D Lightyear™ 1.4, is used to convert the STL files generated during the part preparation and design stages into the build file format required for the operation of the SL 250/50.The entire build process is designed in the form of an assembly using SolidWorks CAD design software. The assembly is first saved as a part file and then converted into an STL file. The final STL file is divided into three distinct parts, as shown in Figure 4, to facilitate the649SL/DW integration. The first STL part is the SL base into which the electronics components are embedded (Figure 4a).

The first STL part is the SL base into which the electronics components are embedded (Figure 4a). The electronic components that are embedded in the SL part include a battery holder, LM 555 timer chip, three resistors, a thermistor, a capacitor, a switch and two LEDs.The 555 timer is one of the most popular and versatile integrated circuits and can be configured as either a monostable multivibrator or as an astable multivibrator (oscillator). It includes 23 transistors, 2 diodes and 16 resistors on a silicon chip installed in an 8-pin mini dual- in-line package. The 555 timer circuit is based on the sequential charging and discharging of the external capacitor and can be used with any power supply in the range of 5-15 volts making it useful in many analog circuits. The simple temperature-sensitive circuit oscillated the LED at a frequency proportional to the temperature sensed by the thermistor.

The simple temperature-sensitive circuit oscillated the LED at a frequency proportional to the temperature sensed by the thermistor. This circuit was selected because it (1) represented a simple circuit combining many typically used electronic components and thus provided a useful demonstration for integrated electronic systems manufacturing applicable to a wide variety of devices, and (2) provided an indication of the parasitic resistances and capacitances introduced by the fabrication process due to its sensitivity to manufacturing variation.In the hybrid operation, traditional SL part building was performed layer by layer with planned interruptions to incorporate the electronic components. The vat was lowered while keeping the build platform location constant to provide access to the sockets and vias for embedding the electronic components.

The vat was lowered while keeping the build platform location constant to provide access to the sockets and vias for embedding the electronic components. Once the components were manually inserted into their respective sockets and their leads were aligned perpendicular to the SL part surface, the vat was re-leveled to the last built layer, thus filling up the sockets and embedding the components with the SL resin. The second STL file was used to cure the resin in the sockets and embed the components followed by a few more build layers to create a flat top surface with matching holes to provide access to the connector pins of the embedded electronic components as shown in Figure 4b. The DW system was used to dispense the conductive ink in the required profile and the connector pin holes are filled with ink to complete the interconnections as shown in Figure 4c. The ink can be cured using either the SL laser or standard heat convection oven at the recommended curing conditions.

The ink can be cured using either the SL laser or standard heat convection oven at the recommended curing conditions. Since this was a preliminary study for embedding functional electronic components and circuits using RP technologies, the heat convection oven curing method was used to cure the ink. However, Medina et al., (2005) previously demonstrated successful curing of the DW ink using the SL laser. In order to enable vertical integration of the electronic components and systems and justify the advantages of this research, optimum laser curing of the ink is the next natural progression in this research for truly automatic fabrication of multi-layer 3D circuitry.The precision 3-axis stages were controlled using LabVIEW®. Registration between the DW tip and the SL part were achieved by locating the dispensing tip on top of the previously described reference point using the stages.

Registration between the DW tip and the SL part were achieved by locating the dispensing tip on top of the previously described reference point using the stages. The dispensing nozzle was controlled using the Z- stage traverse, accurately positioned over the registration point via a digital position readout using the linear encoders on the stages. The Z-axis traverse was adjusted to 4-mils vertical distance so that the required gap between the dispensing tip and the SL part was obtained. These dispensing conditions were chosen based on previous research which provided the optimum dispensing pressure and z distance between the dispensing tip and the SL part for consistent DW traces. After the initial calibration and accurate positioning of the dispensing tip was complete, the DW process was performed using LabVIEW®.

After the initial calibration and accurate positioning of the dispensing tip was complete, the DW process was performed using LabVIEW®. Once the traverse sequence (which a set of650simple motion commands) was activated using LabVIEW®, the dispensing could be controlled to start and stop as and when necessary to complete the entire DW process in one step.After the DW process was complete, the part was then cured in the convection oven to complete the ink curing process. The effect of curing was investigated by testing the resistivity using a MeTex® multimeter (Model # M–3850D). The curing process was repeated until the required resistivity was achieved. The ink was considered cured when the resistivity was within 3% of the previous measurement after each curing cycle.

The curing process was repeated until the required resistivity was achieved. The ink was considered cured when the resistivity was within 3% of the previous measurement after each curing cycle. The SL part was then allowed to cool to room temperature before inserting the battery.Fabrication demonstrationThe complete LM 555 timer circuit was manufactured using DSM Somos® ProtoThermTM resin and the E 1660-136 conductive ink and the hybrid SL/DW system. Figure 5 represents the SL/DW procedure used to manufacture the functional part with embedded electronics and depicts the part at various stages during the build, while still attached to the build platform. The entire process takes place without removing the part from the SL machine platform until the completion of the DW process and thus provides a continuous manufacturing system without the need for re-registration.

The DW stages were controlled using LabVIEW® to replicate the required DW profile and dispense the ink to draw the circuitry. The SL/DW process was initiated by building the SL base with sockets and vias, for inserting the electronic components, using the first set of STL files as shown in Figure 5a. Once the SL part reached a predetermined layer in the STL file, the build was interrupted without raising the platform, the vat was lowered to expose the top surface, and the sockets and vias were cleaned prior to inserting the electronic components. The components were manually inserted into their respective sockets and vias such that the connector pins of the components were slightly below the last built layer and aligned perpendicular to the SL part surface thus allowing the build to continue with regular ZephyrTM blade sweeping. The resin level was re-adjusted and the build was continued to embed the components using the second set of STL files.

The resin level was re-adjusted and the build was continued to embed the components using the second set of STL files. Here, the components were embedded while providing access to the connector pins of the electronic components via precise holes manufactured during the SL build as seen in Figure 5b. The vat was lowered to expose the top surface of the SL part to begin the DW process. LabVIEW® was used to guide the traverse system as well as control the dispensing for the DW interconnects using the 2405 Ultra Dispensing System. The dispensing syringe was directed from a predetermined home position (which allowed normal SL build without any obstructions) to the starting position for the DW process (shown in Figure 5c) using the precision 3-axis traverse stages.

The traverse sequence (which a set of simple motion commands) was initiated, and the dispensing was controlled to start and stop as and when necessary to complete the entire DW circuit in one sequence.After the ink deposition was complete, the functional part was removed from the machine, and cured using a standard heat convection oven. To test the effectiveness of the ink curing, the resistivity of the functional test piece, placed in a heat convection oven at 138 ˚C, was monitored and recorded at 15 min intervals. The curing process was continued until the resistivity value reached the value recommended by the ink manufacturer (0.01 (cid:525)/square/mil), and it was found that curing at 138 °C produced the required resistivity value. The total time for manufacturing the completed part with embedded components and DW interconnections was approximately three and a half hours.

The total time for manufacturing the completed part with embedded components and DW interconnections was approximately three and a half hours. The functional hybrid SL/DW part was further compared to the traditional breadboard LM 555 timer circuit for functionality.651DW start point(a) (b) (c)Figure 5. (a) SL part with sockets; (b) embedded components; (c) DW interconnects.Results and discussionsThe LM 555 circuit manufactured using the hybrid SL/DW machine was analyzed to validate its performance against its equivalent temperature circuit built on a breadboard. The analysis included testing both circuits concurrently at a temperature range of -20°C to 90 °C using a heat convection oven and a freezer. The critical resistances and capacitance values that affected the frequency were measured using a standard multimeter for each circuit.

The critical resistances and capacitance values that affected the frequency were measured using a standard multimeter for each circuit. Both circuits were maintained for 5 minutes at the desired temperatures to approximately achieve thermal equilibrium prior to measuring the resistances and capacitances across the components, which affected the frequency of the LEDs. The circuit drew virtually no power so self-heating during operation was negligible. The resistance and capacitance of the thermistor, resistor (1k(cid:525)) and capacitor on the breadboard circuit were measured by placing standard alligator clips on the leads of the components and measured using the multimeter. For measuring the parameters on the RP circuit, standard wire jumpers were wrapped around the alligator clips.

For measuring the parameters on the RP circuit, standard wire jumpers were wrapped around the alligator clips. The leads of the wire jumpers were inserted into the DW ink lines across the components of the RP circuit, consequently including any parasitic variation that might be induced by the DW ink lines.Table 1.

The thermistor’s properties were confirmed in Table 1, which exhibit an increase in its resistance as temperature decreases and vice versa. The capacitance in the temperature circuit is independent of temperature and varied slightly from its selected value, but is required to determine the frequency of oscillation of the output. The calculated time indicates the period of the blinking LEDs and the differences between the circuits are likely dominated by the component variation rather than the DW ink.12.00Rapid Prototyped CircuitBread Board Circuit10.00) z H8.00(y c n e u q e r F6.004.002.000.0020 -15-10 -5 0 5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90Temperature (°C)Figure 7. Temperature vs. LED frequency comparison14000Rapid Prototyped CircuitBread Board Circuit12000) sm h o (100008000e c n a t s i s e R600040002000020 -15 -10 -5 0 5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90Temperature (°C )Figure 8.

Temperature vs. thermistor resistance comparison653Figure 7 illustrates the comparison of the RP circuit and breadboard circuit under different temperature settings and indicates that temperature is exponentially related to the resistance of the thermistor, as expected.Figure 8 verifies that the temperature is linearly related to frequency of both the RP circuit and the breadboard circuit as predicted at different temperatures. Results from the temperature test conclude that the RP temperature circuit operates similar to the breadboard circuit as little variation in the resistances and capacitances was measured.Conclusions and future workA hybrid SL/DW machine setup was used for manufacturing embedded electronic components in a semi-automated environment with DW interconnects.

The modified SL system was incorporated with build interrupts used to stop and start the SL build while embedding electronic components and using DW interconnects to complete the electrical interconnects. The DW system employed a three-axis translation mechanism that was integrated with the commercial SL machine. The 3-axis DW traverse system and the ink dispensing were controlled using LabVIEW® to replicate the required DW profile and dispense the ink to complete the required circuitry. Comparisons of the RP fabricated electronic circuit with a breadboard circuit indicated that the RP circuit operated essentially the same as the traditional breadboard circuit. It should be noted, however, that this circuit consumed very little power (<10mW), and therefore did not generate any appreciable heat. Circuits with higher power consumption may self-heat due to the reduced thermal conductivity of the system based on the complete encapsulation of the components in the SL resin.

Circuits with higher power consumption may self-heat due to the reduced thermal conductivity of the system based on the complete encapsulation of the components in the SL resin. However, we believe that these challenges can be overcome by embedding novel passive cooling techniques (i.e. heat pipes) and advanced thermal management is the focus of future work.In summary, RP designed circuits can reduce process planning steps enabling cost reduction with reduced assembly due to effective use of material and minimization of waste and effluent. The use of different materials and the ability to build in arbitrary geometries can enable the design and manufacture of compact and visually appealing 3-D circuits. Improved performance by reducing electrical interference between the components and opting to fabricate in several structured functional layers can reduce circuit size and area.

Improved performance by reducing electrical interference between the components and opting to fabricate in several structured functional layers can reduce circuit size and area. By having components embedded within the design, they are protected from outside hazards and weather which increase the overall strength and reliability of the final product. Some of the manufacturing and design issues that can be foreseen during this work include vertical interconnections between components and multiple circuits based on complexity, temperature control during laser curing of ink and its effect on the electronic components, and heat dissipation when using high speed passive components operating at high frequencies.AcknowledgmentsThe research presented here was performed at UTEP in the W.M. Keck Border Biomedical Manufacturing and Engineering Laboratory using equipment purchased through Grant Number 11804 from the W.M. Keck Foundation.

Keck Border Biomedical Manufacturing and Engineering Laboratory using equipment purchased through Grant Number 11804 from the W.M. Keck Foundation. This material is based in part upon work supported by the Texas Advanced Research (Advanced Technology/Technology Development and Transfer) Program under Grant Number 003661-0020-2003. Support was also provided through the Mr. and Mrs. MacIntosh Murchison Chair I in Engineering and through research contract 504004 from Sandia National Laboratories in the Laboratory Directed Research and Development (LDRD) program. Sandia National Laboratories is a multi-program laboratory operated by654Sandia Corporation, a Lockheed Martin Company, for the United States Department of Energy’s National Nuclear Security Administration under contract DE-AC04-94AL85000.References1. D. Begaye, “Optical Transmission on Digital Signals through Rapid Prototyping Fibers”, 2003 National Conference on Undergraduate research, Salt Lake City, Utah.2. P.W.

panding the Applicability of FDM-type Technologies Through Materials DevelopmentD. A. Roberson 1,2,*, C. M. Shemelya 1,3 , E. MacDonald 1,3 and R. B. Wicker 1,41. W. M. Keck Center for 3D Innovation,The University of Texas at El Paso, El Paso, TX 79968 2. Department of Metallurgical and Materials Engineering, The University of Texas at El Paso, El Paso, TX 79968 3. Department of Electrical and Computer Engineering, The University of Texas at El Paso, El Paso, TX 799684. Department of Mechanical Engineering,The University of Texas at El Paso, El Paso, TX 79968Corresponding Author: D. A. Roberson, droberson@utep.eduAbstractCurrently, the most common form of additive manufacturing is material extrusion 3D printing (ME3DP) based on fused deposition modeling (FDM®) technology which relies upon a thermoplastic monofilament as a base material for the fabrication of three dimensional objects.

The dependence on thermoplastics as a feedstock by ME3DP platforms limits the applicability of this additive manufacturing method. A clear-cut path towards greater applicability is the introduction of novel materials with diverse physical properties which maintain compatibility with 3D printing platforms based on FDM® technology. The work in this paper presents efforts in the development of polymer matrix composites (PMC)s and polymer blends based on acrylonitrile butadiene styrene (ABS) and polycarbonate (PC), two thermoplastic materials commonly used by FDM®-type platforms. Mechanical testing and fractography via scanning electron microscopy (SEM) were the two main metrics used to characterize these new material systems.

Mechanical testing and fractography via scanning electron microscopy (SEM) were the two main metrics used to characterize these new material systems. Overcoming barriers to the manufacturing of these novel 3D-printable materials systems is also presented.Key Words: 3D Printing; Materials Characterization; Polymer Matrix CompositesIntroductionThe evolution of additive manufacturing (AM) from a tool used to rapidly create three dimensional models into to a fully developed technique capable of fabricating multifunctional devices has undergone significant advancements in recent years. One driving force of significant innovation has been the hybridization of AM with other manufacturing techniques such as direct write (DW) where conductive and insulating materials are deposited enabling the creation of “structural” electronics.

There are several examples of these novel electronic devices where DW was combined with stereolithography (SLA); pushing the forefront of what is possible using AM [1-5].While such achievements help to advance the field of AM to greater heights in applicability and benefit to society, 3D printing platforms must rely, for the most part, on “off the shelf” material systems. For example, there is a large body of work pertaining tocharacterization of titanium alloys (typically Ti 6Al 4V) as used in electron beam melting (EBM) and selective laser melting (SLM) [6-9]. So too are there many instances of the characterization of acrylonitrile butadiene styrene (ABS) for use in FDM applications. However, by allowing a given platform access to materials with a diverse set of physical properties, the possibility to fabricate all-3D-printed, multi-functional structures becomes reality.

However, by allowing a given platform access to materials with a diverse set of physical properties, the possibility to fabricate all-3D-printed, multi-functional structures becomes reality. One such example is the Objet material jetting AM platform which can print either epoxy or elastomeric materials. This enables the printing of objects with both flexible and rigid sections, and has led to the ability to fabricate actuators and other objects which can take advantage of rigid and flexible members such as mechanical actuators among others [10, 11].The work presented in this paper demonstrates the development of application-specific material systems meant for use in material extrusion 3D printing (ME3DP) platforms through the creation of novel polymer matrix composites (PMC)s and polymer blends (PB)s.

The new material systems with three intended applications: 1) the application of ME3DP in electromagnetic and electromechanical uses; 2) the application of ME3DP in austere environments; and 3) the application of ME3DP itself. The third goal was geared towards the development of materials which mitigate issues associated with ME3DP such as build orientation-related mechanical property anisotropy— an issue documented in the use of nearly every AM platform [12-24]— as well as aspects other such as surface finish.The strategy employed by our group, to further the applicability of ME3DP, has been the creation of PMCs and PBs based upon a known printable material such as ABS or polycarbonate (PC).

The common theme found between the fabrication of PMCs and PBs is the taking advantage of the physical properties of two or more materials, which have been blended together in some way.In general, the equation describing a composite is typically given in terms of yieldstrength:,where σc is the yield strength of the composite, Vfm is the volume fraction of the matrix( in our case ABS or a similar thermoplastic), Vfr, is the volume fraction of the reinforcing agent, σm is the ultimate tensile strength of the matrix and σr is the ultimate tensile strength of the reinforcing agent.While there are several parameters of a material which can be tailored by polymer blending, an example of blending two polymers in order to manipulate glass transition temperature (Tg) comes in the form of the Fox equation [25]:,where x1 and x2 are the weight fraction of the individual polymers and Tg1 and Tg2 represent the glass transition temperatures of the two polymers in a blend.

An example of a polymer blend used in FDM is Ultem 9085 which is a blend of polyetherimide (Tg = ~216 °C) and PC (Tg = ~147 °C). In this example, Ultem 9085 has a lower Tg (~186 °C) compared to(1)(2)polyetherimide alone due to the addition of PC. As will be seen in this work various representations of the rule of mixtures such as equations 1 and 2 are the lynchpin to increasing the diversity in physical properties of 3D printable materials.Experimental ProcedureAll PMCs and blends were fabricated through the use of a Dr. Collin twin screw extruder / compounder (Model ZK 25T, Dr. Collin GmbH, Ebersberg, Germany) which was designed to fabricate a monofilament 1.75mm in diameter. Each material system necessitated specific machine properties which can be found in previously reported work [18, 26]. Mechanical testing was carried out through the use of an Instron® 5866 tensile test machine (Instron, Norwood, MA).

Mechanical testing was carried out through the use of an Instron® 5866 tensile test machine (Instron, Norwood, MA). Tensile test specimens were printed using a MakerBot Replicator (MakerBot Industries, Brooklyn, NY) following dimensions specified in the ASTM D638-10 based on the Type V parameters [27]. Fractography was performed on the fracture surfaces of tested specimens by analyzing micrographs obtained from a scanning electron microscope (SEM). The SEM used in this study was a Hitachi TM-1000 Tabletop SEM equipped with a backscatter electron detector and operating with an accelerating potential of 15 kV (Hitachi High-Technologies Europe GmbH, Germany).

As we were examining polymeric specimens, a preliminary sputter coating with an Au/Pd alloy was necessary and carried out through the use of a Gatan coating system (Model 682, Gatan, Inc., Pleasanton, CA).Electromagnetic and Electromechanical ApplicationsAdditive Manufacturing has proven to be a key enabler for the creation of metamaterial devices due to the ability to fabricate complex anisotropic and spatially variant geometries [28, 29] as seen in Figure 1. The major drawback to the use of AM in this application is the the electromagnetic (EM) properties of the dielectric feedstock material, namely constant— which for most is ~2 thermoplastics. Analogous to increasing the yield strength of a polymer through the addition of carbon or glass fibers, the dielectric constant of a polymer can be increased through the addition of materials composites (metal oxides for example) with a greater relative permittivity.

Many equations have been developed to model the dielectric constant of a mixture; one being the Maxwell Garnett equation [30]:to 3Figure 1. 3D printed anisotropic metamaterial [28].,where is the permittivity of the mixture, is the permittivity of the matrix, is the permittivity of the additive material, and is the volume fraction of the additive material.(3)Work by our group has dealt with the blending of ABS and PC with TiO2 (ε=50) in the development of printable materials geared for EM applications. One problem encountered was achieving dispersion of the additive within the matrix. We found it necessary to functionalize TiO2 in a process involving a Silane agent (purchased from Sigma-Aldrich) geared for either PC or ABS which resulted in smaller agglomerations and better general dispersion (Fig.

2).Another application which stands to befit from the availability of printable materials with a wider range of physical properties is the 3DP printing of an electric motor (Fig. 3) as demonstrated by Aguilera et al. [31]. Were the core of the motor to be printed from ferromagnetic materials, the power of the motor would be increased. En route to achieving this goal, we have developed printable ferromagnetic composites based on either PC or ABS (Fig. 3)Figure 2. Functionalizing TiO2 prior to compounding with ABS led to better dispersion within the polymeric matrix.Figure 3. a) The process of fabricating a 3D printed motor. From reference [31].

Functionalizing TiO2 prior to compounding with ABS led to better dispersion within the polymeric matrix.Figure 3. a) The process of fabricating a 3D printed motor. From reference [31]. b) 3D-printable ferromagnetic material which could be used to enhance the motor’s performance.Applications in Austere EnvironmentsIn order for space-based and remote research outposts to effectively utilize ME3DP, the implementation of a closed-loop manufacturing scheme is needed where printed parts can be reprocessed into a 3D printer-compatible monofilament. This practice would enable more 3Dprinted structures from the same amount of base materials and allow for the reprocessing of short lifespan or misprinted parts; requiring less material to be transported to a given production site.

However, a concern for such a material recycling process is the thermomechanical degradation which will result in decreased mechanical strength [32].There has also been significant investigation of the effect of additives on the thermomechanical degradation of ABS. Here the additives investigated were ZnO nanorods, TiO2 nanoparticles, and a palygorskite organo-nanoclay pigment marketed as MayaCrom® blue. ABS-based PMCs created from each of the three additives were prepared and printed as tensile test specimens. The specimens were tested, reprocessed into a monofilament, and then used to print another second set of tensile specimens. The process was then repeated for two recycling cycles. As can be seen in Fig. 4, the addition of 2% by weight ZnO nanorods demonstrated the most promise in the reduction of thermomechanical degradation due to reprocessing.Figure 4. Results of thermomechanical degradation testing on ABS-based composites.

Results of thermomechanical degradation testing on ABS-based composites. Here ZnO nanorods demonstrated the most promise in the mitigation of thermomechanical degradation.Of particular interest to space-based manufacturing is the ability to print materials with radiation shield capability. One metric to test this ability is testing of the transmittance of x-rays through a material. The intensity of x-rays which pass (or are transmitted) through a material of a given thickness, x, can be calculated based on the following equation [33]:(),where is the intensity of the x-rays before passing through the material and ρ is the density of is known as the mass absorption coefficient and is pertinent to our the material.

The parameterdevelopment of application-specific PMCs as it can be manipulated through mixing similar to the other physical parameters discussed here as described by the equation [33]:(4)()()(),where the subscript m denotes mixture, and are the mass fractions of substance 1 and substance 2. It should also be noted that density plays a large role in the blocking of x-ray transmission as indicated by equation 4. In our case we chose to improve the x-ray impeding capability of PC through the addition of tungsten powder and a simple model to describe the density of our composite can be expressed by the equation:,which essentially demonstrates the manipulation of a PMC through the addition of dense material. Indeed it has been shown that even small amounts of tungsten can have a profound effect on the ability to block x-ray radiation transmission.

Indeed it has been shown that even small amounts of tungsten can have a profound effect on the ability to block x-ray radiation transmission. Figure 5 shows the difference in x-ray transmission between 3D printed plates fabricated from PC and PC/ W PMCs loaded with 1%, 3% and 5% by weight tungsten, respectively.Figure 5. a)The ability to tailor the amount of x-ray transmission based on tungsten loading and b) 3D printed impact test specimens made from the same materials.Materials for the Application of 3D PrintingAs mentioned before; currently, the vast majority of manufacturing based on 3D printing relies on traditional or “off the shelf” material systems.

While there are examples of the development of novel materials geared towards the 3D printed fabrication of novel devices, there remains a need for the development of materials geared specifically for 3D printing.The development of 3D-printer-specific material systems should be directed towards those that mitigate issues associated with a given 3DP process. As mentioned before, an issue encountered by nearly every AM process is mechanical anisotropy based on build orientation. Initial characterization of the effect on additives on build orientation anisotropy was presented in Torrado Perez et al.

Initial characterization of the effect on additives on build orientation anisotropy was presented in Torrado Perez et al. [18] where a blend of ABS and 5% by weight styrene ethylene butadiene(5)(6)styrene (SEBS) was shown to decrease the difference in ultimate tensile strength (UTS) between samples printed in the XYZ and ZXY directions.Further development of binary and ternary ABS-based PBs has been presented by Rocha et al.[26] where it was shown that a ternary blend of ABS, ultra-high molecular weight polyethylene (UHMWPE) and SEBS in a by weight ratio of 75:25:10 was shown to be able to print smoother inclined planes as compared to ABS alone. Analysis via scanning electron microscopy (SEM) revealed that the rheology of this ternary PB led to an intermingling of the print rasters as seen in Figure 6. The rheological differences altered the deposition characteristics as compared to ABS and allowed for the printing of smoother planes.Figure 6.

The rheological differences altered the deposition characteristics as compared to ABS and allowed for the printing of smoother planes.Figure 6. SEM images of cross sections of a) ABS, b) ABS:SEBS 50:50 blend and c) ABS:UHMWPE:SEBS 75:25:10. Note the differences in print rasters. From [26] d) Corresponding surface roughness data from a test piece described in Rocha et al. [26] showing the ability of these novel blends to print smoother inclined planes.The same blend has also proven to be an enabler for the improvement of anisotropy by decreasing the difference in UTS between samples fabricated in the XYZ and ZXY print directions [34] as is demonstrated in Figure 7. The reason for this decrease was shown to be due to the intermingling between print rasters as seen by the cross sections in Figures 6 and 7 and the fracture surfaces in Figure 8.

The reason for this decrease was shown to be due to the intermingling between print rasters as seen by the cross sections in Figures 6 and 7 and the fracture surfaces in Figure 8. While this blend does experience a dramatic decrease in UTS, as compared to ABS, it stands as a step towards the development of a material system geared specifically at mitigating an issue inherent to AM.Figure 7. The rheological differences of the ternary blend as compared to ABS obscure the print rasters leading to a decrease in build orientation-caused mechanical property anisotropy [26, 34].Figure 8. The intermingling in print rasters is also apparent on the fracture surfaces of tensile test specimens printed in different orientations from [34].ConclusionsThe work presented here demonstrates research efforts geared towards the development of 3D printable materials which are intended for specific applications involving material extrusion 3D printing.

Through the development of novel polymer matrix composites and polymer blends, new materials have been fabricated which can be applied to: 1) 3D printing of electromagnetic and electromechanical components; 2) 3D printing in austere environments; and 3) material extrusion 3D printing. The strategy of utilizing known printable base materials has aided in maintaining printing compatibility with material extrusion 3D printing platforms.Materials characterization efforts based on mechanical testing and SEM microanalysis have provided insight into the dispersion of additives, the rheological behavior of new material systems, and the fracture morphology. This information is critical in building a knowledge base for the development of new thermoplastic material systems for 3D printing.

This information is critical in building a knowledge base for the development of new thermoplastic material systems for 3D printing. While the advances towards the creation of all-3D printed electronic, electromechanical, and electromagnetic devices will continue, the development of novel material systems remains a critical enabler for the future of additive manufacturing.AcknowledgementsThe work performed here occurred in the W.M. Keck Center for 3D Innovation and the Department of Metallurgical and Materials Engineering at The University of Texas at El Paso. The authors are grateful for the contributions of Joel English, Angel Torrado Perez, Carmen Rocha and Lauro Barbieri. This project was supported by a grant from the Intelligence Community Postdoctoral Research Fellowship Program through funding from the Office of the Director of National Intelligence under grant no. 2012-12071000005.

2012-12071000005. All statements of fact, opinion, or analysis expressed are those of the author and do not reflect the official positions or views of the Intelligence Community or any other U.S. Government agency. Nothing in the contents should be construed as asserting or implying U.S. Government authentication of information or Intelligence Community endorsement of the author’s views. The authors are alsograteful for additional support which came from the National Aeronautics and Space Administration (NASA) under grant no. 282002-8784.References[1]A. J. Lopes, E. MacDonald, and R. B. Wicker, “Integrating stereolithography and direct print technologies for 3D structural electronics fabrication,” Rapid Prototyping Journal, vol. 18, no. 2, pp. 129–143, 2012.[2]R. B. Wicker and E. W. MacDonald, “Multi-material, multi-technologystereolithography,” Virtual and Physical Prototyping, vol. 7, no. 3, pp.

An analytical pulsed laser thermal model establishes the modeling procedure. The framework inverts a powder bed’s single pulse temperature response from experimental spatiotemporal Short-Wave Infrared (SWIR) camera data. Superimposing this response along a scan path simulates multi-pulse LPBF. Results show the experimentally informed superposition model rapidly and accurately predicts a layer’s temperature history. The model has applications in correction of thermally driven LPBF errors and in-situ part qualification.IntroductionLaser Powder Bed Fusion (LPBF) fabricates high resolution, complex, metal parts layer- by-layer. Part geometry variances and changing laser scan pathing drive thermal differences within layers that produce defects [1]. Analytical [2], numerical [3], and high fidelity [4] thermal models aim to understand LPBF’s underlying physics causing those variations with significant computational expense.

Analytical [2], numerical [3], and high fidelity [4] thermal models aim to understand LPBF’s underlying physics causing those variations with significant computational expense. Controls-oriented superposition thermal models quickly predict a layer’s temperature history. The superposition approach’s linear assumptions reduce a layer’s computation time from days, or weeks, to minutes, making it feasible for application in real-time process correction.Moran et al. combined Rosenthal’s moving point heat source solution with FEA to perform superposition simulations of a laser scanning Ti-6Al-4V [5]. Moran et al. found their model provides results accurate to FEA only simulations with a significant decrease in computation time. The computation time reduction makes large scale simulations practical. Schwalbach et al. used Green’s function methodology to develop a thermal model for LPBF with a continuous wave (CW) laser [6].

The computation time reduction makes large scale simulations practical. Schwalbach et al. used Green’s function methodology to develop a thermal model for LPBF with a continuous wave (CW) laser [6]. Schwalbach et al.’s model approximates a scanning CW laser by superimposing the temperature response from distinct heat sources seeded along the raster path. Schwalbach et al. demonstrated their model’s temperature results agree with analytical solutions, calibration provides accurate melt pool dimension predictions, and the approach indicates thermal history spatial variations.1326Temperature and state dependent thermal properties, latent heat effects from melting and solidification, and material vaporization make the LPBF process highly non-linear. This limits analytical superposition models to qualitatively predicting part geometry-scan path interaction effects.

This limits analytical superposition models to qualitatively predicting part geometry-scan path interaction effects. A superposition model informed by thermal camera data would provide quantitative predictions with LPBF’s physics embedded, but this requires the experimental measurements to behave linearly. The quantitative analysis may improve decision making capability, and baseline in-situ measurements to assist part qualification.This paper presents an LPBF superposition thermal model constructed with Short-Wave Infrared (SWIR) thermal camera measurements. A theoretical superposition model’s derivation for pulsed laser LPBF reveals the methodology for experiment-based modeling. The superposition modeling process gauges the SWIR camera measurement’s linearity.

The superposition modeling process gauges the SWIR camera measurement’s linearity. Processing simulation and experimental data with an in-situ framework evaluates the superposition model’s ability to predict real thermal history variances.Model DevelopmentThe theoretical pulsed laser superposition model’s derivation follows Schwalbach et al.’swork for CW systems [6]. The pulsed laser’s volumetric heat source model,(),,,qxyzt, is [7]eppzRPxyqxyztttδπωωδ()()()222221,,,exp2exp1−+=−−−−,where(),,xyzare the spatial coordinates, t is the time, R is the reflectance, P is the laser power,δp is the laser’s optical penetration depth, ω is the beam waist, Θ is the Heaviside function, and te , a single laser pulse produces in a semi- is the exposure time.

The basis function copies offset spatially according to the laser scan path and temporally by the pulse period. The temperature for a multi-pulse simulation at z = 0 is()()01,,,==+KkkTxytTGrt,where Gk is the kth laser pulse’s basis function. The kth pulse’s radius, r, to a given (x,y) and its relative time,t, are220,0,0,=−+−=−kkkrxxyyttt()(),where()0,0,,kkxyare its exposure coordinates and0,ktis when exposure begins.Figure 1 demonstrates superposition for six laser pulses (P = 200 W, te = 70 µs, ω = 70 µm, δp = 60 µm) striking 304L stainless steel (k = 15 W/m·K, ρ = 7800 kg/m3, cp = 500 J/kg·K). Figure 1 (a) contains the pulses’ basis function profile plots (purple to red curves) at their relative times, offset in space by the laser point-to-point distance (60 µm). The profiles demonstrate the basis function’s temperature magnitude decreases, and its waist increases after the exposure time as heat conducts away from the exposure location.

The camera samples at 3345.8 Hz with 130 µm/pixel and 135 µm/pixel x and y instantaneous field of views, respectively (see [8] for further details). The raw camera data corrects to temperature using a procedure like discussed in [9]. Figure 2 (a) shows a temperature corrected image with the laser’s exposure points (white dots) and parameters annotated. The laser scans with a 0° rotation angle, a 60 µm point-to-point distance, dp, and an 85 µm hatch spacing, dh. Each image supplies an A sub-matrix and a ’s length, after 10 ms, or beyond a radius of 0.78 the procedure builds A sub-matrices by assuming has resolutions of Δt = 150 µs and Δr = 130 µm. Data from 500 mm. Also, the process assumes consecutive thermal images assemble A and . Multiple images provide a full rank A and reduceGTsub-vector. Since A’s columns must equal0G→GGTto invertG’s noise.

This process produces a thermally based data map for layers. Thermal features have application in part property correlations and in-situ defect detection. A layer’s thermal feature map also clearly show how the laser’s scan path affects the thermal history. Two common thermal features are the maximum temperature [10] and the time above threshold [11]. Figure 4 evaluates the superposition simulation’s predictions of those thermal features. Figure 4 contains experimental and simulated thermal feature profile plots and maps. The maximum temperature, Tmax, predictions in Fig. 4 (a) and (c) match the experiments in the interior but deviate at the layer’s edges. This suggests a linearity reduction where the laser corners. The entire layer’s maximum temperature RMSE is 177 K. Figure 4 (b) and (d) illustrate the simulation successfully predicts the experimental time above threshold, τ. Time above threshold experimental deviations occur at some locations due to balling.

The successful basis function inversion demonstrates the SWIR camera temperature data’s linearity for nominal manufacturing. The superposition approach simulates a simple 5×5 mm2 layer in 4 minutes. The simulation’s analyzed time series data fits the experiment’s data with a maximum 174 K RMSE. Simulation thermal features track experimental results with a 177 K maximum temperature RMSE and a 1.2 ms time above threshold RMSE. The pulsed laser experimental superposition framework should adapt to CW LPBF systems since thermal cameras have finite sampling rates. The superposition model’s applications include laser scan path effects prediction to baseline experimental data in part qualification, and feedforward controls. Future work will explore experimental superposition for various LPBF manufacturing cases and evaluate the ability to flag defects in parts.AcknowledgementThis work was funded by Honeywell Federal Manufacturing & Technologies under Contract No.

DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.13322015T m a x ‒ T 0[10K · 1 0 2 ]1 mm10τ5[m s ]0References[1][2][3][4]Druzgalski CL, Ashby A, Guss G, King WE, Roehling TT, Matthews MJ. Process optimization of complex geometries using feed forward control for laser powder bed fusion additive manufacturing. Addit Manuf 2020;34:101169. https://doi.org/10.1016/j.addma.2020.101169. Ning J, Sievers DE, Garmestani H, Liang SY. Analytical modeling of part porosity in metal additive manufacturing. Int J Mech Sci 2020;172:105428. https://doi.org/10.1016/j.ijmecsci.2020.105428. Bruna-Rosso C, Demir AG, Previtali B.

viewed PaperEXPERIMENTAL ANALYSIS AND OPTIMIZATION OF GAS FLOW IN AN OPEN-ARCHITECTURE METAL L-PBF SYSTEMMagnus Bolt Kjer∗1, Moritz Reinhard Ludwig Zwicker1, Venkata Karthik Nadimpalli1, Sebastian Aagaard Andersen1 and David Bue Pedersen1 1Department of Mechanical and Civil Engineering, Technical University of Denmark Kgs. Lyngby, Denmark ∗Corresponding author: mbokj@mek.dtu.dkAbstractA sensor platform for experimental in-situ characterization of metal laser powder bed fusion (L- PBF) is developed and used to analyze the gas flow across the powder bed. The gas flow in metal L-PBF is crucial to attaining transient and spatially consistent results. The gas flow removes spatter from the welding process as well as the plume consisting of metal condensate. Defects like porosity will result from these process byproducts not being removed. A CNC gantry with a flow sensor mounted measured the gas velocity across the process area.

Defects like porosity will result from these process byproducts not being removed. A CNC gantry with a flow sensor mounted measured the gas velocity across the process area. The automated measurements were analyzed and used to improve the gas flow by changing the inlet geometry. The method can be used as a supplement to computational fluid dynamics, as well as a calibration tool for simulations, to improve the efforts toward better models of the metal L-PBF process.IntroductionThe work relates to the gas flow of laser powder bed fusion (L-PBF). L-PBF is a powder-based additive manufacturing process, in which a thin layer of powder is selectively fused by a high- power laser. A three-dimensional object can thus be built by subsequent additions of layers and laser fusion.

A three-dimensional object can thus be built by subsequent additions of layers and laser fusion. A L-PBF system can be decomposed into several sub-systems, one of which is the gas flow sub-system - a sub-system which is generally under-appreciated in literature.This paper proposes an experimental platform for characterizing gas flow in metal L-PBF - enabling detailed spatial mapping of the gas flow conditions. The inert gas cross-flow in metal L-PBF is an often overlooked yet crucial sub-system, which is vital in regards to process stabil- ity[1]. Computational Fluid Dynamics (CFD) simulations in the context of metal L-PBF are often described without direct experimental verification. As such, measuring the flow velocity across the powder bed is needed to understand the local flow conditions, and can be used to calibrate and validate said CFD simulations towards better approximations of the process.

If done in an auto- mated manner, a dense collection of measurements can form the entire field[2]. Installing such an automated system requires machine access - and more specifically machine freedom, freedom to alter the machine physically as well as in software. The experimental approach can also be done to quickly iterate the gas flow geometry without the need for updating and solving a complex CFD model. The experimental method also takes the entire system into account - that being e.g. actual1321blower characteristics, filter, piping, etc.The byproducts from the welding process must be removed from the processing area for several reasons. For one, the plume from the melt pool will block the incident laser power. This will cause a decrease in power, and cause instability of the melt pool as this cloud-like effect fluctuates[3]. Secondly, the larger particles which constitute the spatter must be blown away to avoid landing upon non-welded powder.

Secondly, the larger particles which constitute the spatter must be blown away to avoid landing upon non-welded powder. A balance must be found between sufficient flow to maintain a stable process while not disturbing the powder bed[4]. Furthermore, the gas flow performs a supportive function as the flow mitigates the buildup of condensate on the optical and mechanical components.The major effect is typically seen through component density, either due to lack of fusion or balling defects [1, 4–8]. The level of porosity has been shown to correlate with regions across the powder bed with sub-optimal flow conditions [9], meaning the spatial distribution of the gas flow velocity across the powder bed is crucial. Even mechanical properties have been linked to gas flow variations[10].

Even mechanical properties have been linked to gas flow variations[10]. However, many investigations are limited to measuring the overall gas input into the processing area[7] or the inlet of the process chamber[5], sometimes due to outside factors like machine access, warranty concerns, etc. The open-architecture nature of the self-build system enable such alterations and additions - compared to conventional commercial systems - like a custom sensor CNC gantry. This allows for a full and proper mapping of the gas velocity.Gas flow subsystemThe system that is subject to the experiments is an open-architecture metal L-PBF machine engi- neered at the Technical University of Denmark (DTU). The cross-flow subsystem is a closed system and is driven by a side-channel blower1, capable of a flow rate of 180 m3 h and a compression pres- sure of 29kPa. The speed of the blower motor is governed by a variable frequency drive (VFD).

The speed of the blower motor is governed by a variable frequency drive (VFD). The frequency on the VFD is in turn controlled by a unified systems controller also engineered at DTU[11].A bottom inlet duct begins to widen the flow. A top inlet duct bends the flow, where specific geometry shapes the flow and minimizes the turbulence of the open jet stream as in e.g. a wind tunnel[12]. After exiting through the outlet duct, the gas is filtered to ensure a clean process.

a wind tunnel[12]. After exiting through the outlet duct, the gas is filtered to ensure a clean process. While the beneficial effects of a structure consisting of small tubes - like a honeycomb pattern - are well- founded, the actual measured impact on the flow behavior was pursued using the above-mentioned method.The gas flow must be 1) of sufficient speed, at least 2.5-3m/s; 2) be uniform across the powder in order to achieve spatially consistent parts [9]; 3) and be sufficiently low in velocity at the powder bed (z=0), thus leaving the powder undisturbed[5].Experimental method and setupTo analyze the quantity and quality of the gas flow, a 3-axis CNC gantry was engineered in order to move the flow probe in-situ, Figure 2a. An AVR ATmega 8-bit 16 MHz microcontroller was used to control the CNC.1Also known as a ’ring blower’ or ’vortex pump’1322Figure 1: Gas flow subsystemIn Figure 2b, the modular inlet design is shown.

The geometry was able to be changed in order to rapidly compare the effect of the front attachments. The prototypes were printed using FDM printers.An EE671 Miniature Air Flow Probe was used to measure the flow velocity with a range of 0-15 m s and analog output of 0-10V. The flow probe has a stated response time of ts = 4 seconds. 3 seconds of data were analyzed at 100Hz after waiting 4 seconds, i.e. a total dwell time at each position of 7 seconds.While the EE671 is an inexpensive sensor, it was found to be accurate in calibration. The sensor was positioned in the opening of a wind tunnel after a flow restriction of known areas A1 and A2. The air velocity was calculated by measuring the static pressure drop, p1 and p2, across the restriction using a micromanometer.

The air velocity was calculated by measuring the static pressure drop, p1 and p2, across the restriction using a micromanometer. The velocity was found using The Bernoulli Equation,p1 +1 2ρv21 = p2 +1 2ρv2 2and the Equation of Continuity (assuming an incompressible fluid),v1A1 = v2A2v1 is the velocity upstream, and v2 the velocity downstream at the probe. ρ is the density of thefluid.A National Instruments 9215 data acquisition module was used to collect the 0-10V analog sig-nal. The frequency of the blower motor was kept constant at 35Hz throughout the experiments.Generating the measurement locations was done using a Python script. The sampling locations were placed on a series of concentric circles as the Ø250mm powder bed of the L-PBF system is round. This was done for each z-plane to be measured.

The sampling locations were placed on a series of concentric circles as the Ø250mm powder bed of the L-PBF system is round. This was done for each z-plane to be measured. Each point in Figure 3 has a corresponding1323(1)(2)a) Sketch of gantryb) Exploded view of modular inletFigure 2set of (x,y,z) coordinates, each plane having 60 measurement positions per plane, entailing an ap- proximate spacing of 25mm in the x-y-planes. A total of 9 planes with a z-spacing of 10mm were used. The center of the powder bed top surface was designated as the origin po(0, 0, 0) of the flow probe.Figure 3: Sampling pattern of the gas flow characterization in relation to the build plate coordinate system.1324Experimental procedureFirstly, the experiment was made without flow geometry attachments to form a baseline for the geometrical investigation.

A straight extension geometry was installed with thin walls for structural stability, and thereafter the same straight duct but with an internal honeycomb structure. The aspect ratio was 7:1, measuring 35mm long and 5mm wide.The different geometries were inter-compared, and the effect of the various geometries on the gas flow were analyzed. Figure 4 shows the honeycomb inlet guiding the gas across the powder bed, interacting with the flow probe, before exiting through the gas outlet port.Figure 4: Photo of inside the L-PBF machine with the honeycomb prototype and gas outlet along with flow probe.ResultsThe baseline geometry results show a wide cone-shaped flow with unacceptably high flow speeds at the powder bed (z=0), seen in Figure 5a, causing the powder to be blown away. The flow was also asymmetrical and unstable, indicating turbulence. Comparing the three geometries is helped by looking at the data from the XZ-plane in Figure 6.

The flow was also asymmetrical and unstable, indicating turbulence. Comparing the three geometries is helped by looking at the data from the XZ-plane in Figure 6. In this plane, the baseline is clearly resulting in a wide flow, reaching 2 ms nearly the entire powder bed, at z=0.The straight duct without honeycomb showed similar results to the baseline, seen in Figure 5b. The maximum velocities increased by nearly 1m/s, which in this case is unproductive. The inlet is now closer to the powder bed - meaning the start of the free jet is nearer the powder bed. The inadequate laminarity and increased velocity results in excessively high velocities at the powder bed level (z=0). The flow is now slightly more centered, but still biased to one side. The straight duct was found to barely be an improvement when compared to nothing.

The flow is now slightly more centered, but still biased to one side. The straight duct was found to barely be an improvement when compared to nothing. Also from the XZ-plane in Figure 6, the straight duct without honeycomb is not performing better, except the flow is increasing in velocity.Figure 5c shows the results after adding internal honeycombs to the flow geometry. Now, the powder-bed-plane z=0 is acceptable halfway across the powder bed, solvable by minor changes to the geometry or by decreasing the blower flow rate. The maximum flow is now centered, and sufficient velocities are now presents across the entire z=40mm and z=50mm plane. The YZ-plane (y=0) in particular (Figure 6) present how the Honeycomb geometry shapes the flow as intended1325and results in a profile that overshoots the powder bed. The clear distinguishable velocity lines indicate a smooth and laminar flow. The laminar flow ensures, as stated above, that the powder bed is left intact.

The clear distinguishable velocity lines indicate a smooth and laminar flow. The laminar flow ensures, as stated above, that the powder bed is left intact. In addition, the plume is moved in the preferred direction - that being toward the outlet and to the filter.Figure 5: Stacked XY planes of the geometries1326Figure 6: XZ vertical plane of the three geometriesThe YZ-plane (x=0) in Figure 7 clearly visualize the asymmetrical properties in the flow in the y-direction for the baseline and No Honeycomb geometries, respectively. Repeating the experi- ments resulted in identical plots and bias towards the positive Y-direction. For baseline and No Honeycomb, having 2-4 m s is unusable. While the Honeycomb performs significantly better, there is still room for improvement, including the high flow speeds in the z=0 plane for y=70mm and above.

While the Honeycomb performs significantly better, there is still room for improvement, including the high flow speeds in the z=0 plane for y=70mm and above. The velocities of approximately 2 ms at the powder bed are unacceptable.1327Figure 7: YZ vertical plane of the three geometriesDiscussion and ConclusionThe main intention was to develop a spatial sensing platform from which to study the L-PBF system (and process) in situ. The sensing platform in combination with the goal of further optimizing the gas flow in the machine led to an understanding of the importance of the geometries in the gas flow subsystem. While further work is needed, the method was found to enable quick iterations and fast implementations of new designs.The honeycomb structure (or equivalent) was found to be crucial in not only making the flow more laminar but also in evenly distributing the flow across the powder bed.

While the importance of honeycomb was expected[12], the degree to which it influences the flow is striking - it should be noted that the same effect should be possible with a staggered hole pattern [10].The measured decrease in velocity across the powder bed in the x-direction is expected and seen repeatedly in the spatial mapping of gas flow in L-PBF[2, 9, 10], and is acceptable as long as the velocities are sufficient. To figure out what is sufficient requires further work to couple the detailed spatial mapping to printed components. Also, different alloys and energy densities result in different spatter and plume formation behavior and therefore might mean a disparity in gas flow requirements.

Also, different alloys and energy densities result in different spatter and plume formation behavior and therefore might mean a disparity in gas flow requirements. This will increase the number of variables needed to be considered when discussing a truly optimized gas flow.The plots along the XZ-plane indicate that, using internal honeycombs, the inlet can be moved1328even closer to the powder bed with minor geometrical changes and possibly a change of flow rate. The height of the flow relative to the powder bed is surely critical, and compared to commercial systems the design presented in this paper is further away from the powder bed[10]. The plume, while violent and fast[13], can and must be removed from the incoming laser as quickly as possi- ble[8].

The plume, while violent and fast[13], can and must be removed from the incoming laser as quickly as possi- ble[8]. Lowering the gas flow is likely beneficial in that regard.The sensing platform holds potential for other sensors than flow probes, e.g., oxygen, tempera- ture, or various cameras could be utilized for a detailed in-situ mapping of the process, even while manufacturing. A line camera has been shown to reveal details about the powder bed[14].While the presented method was an effective method in which to iterate on the gas flow design, the method is also intended for collecting data used for calibrating simulations, with the intention of improving the models further.References1. Ladewig, A., Schlick, G., Fisser, M., Schulze, V. & Glatzel, U. Influence of the shielding gas flow on the removal of process by-products in the selective laser melting process. Additive Manufacturing 10, 1–9. ISSN: 22148604 (Apr. 2016).2. Schniedenharn, M., Wiedemann, F.

viewed PaperExperimental Analysis of Functionally Graded Materials produced by Fused Filament FabricationSuhas Alkunte1, Dr. Ismail Fidan 2, Seymur Hasanov3 1Department of Mechanical Engineering Tennessee Tech University, Cookeville, TN 38505 2Department of Manufacturing and Engineering Technology Tennessee Tech University, Cookeville, TN 38505 3College of Professional Studies The University of Alabama in Huntsville, Huntsville, AL 35899AbstractMulti-material additive manufacturing has grabbed tremendous attention in the research community. In this investigation, a multi-material single extrusion system was used to fabricate the combination of chopped carbon fiber reinforced Polyethylene Terephthalate Glycol (CF- PETG) and Thermoplastic Polyurethane (TPU) materials with gradient transition for a more robust material interface.

Various patterns such as the 20, 40, 60, and 80% by volume blend of CF-PETG and TPU materials are designed, printed, and analyzed to understand their tensile and fatigue behaviors. Tensile–tensile fatigue tests with a stress ratio of 0.1 were performed on each specimen at 80% of UTS. The characterization of functionally gradient material interface and direct transition patterns were conducted for comparison. The results showed that gradient change in material concentrations from soft to hard material has significantly enhanced the interface strength.Keywords: Additive Manufacturing, Functionally Graded Materials (FGM), Tensile, Fatigue.IntroductionFunctionally Graded Additive Manufacturing (FGAM) is layer by layer manufacturing process. It allows for the simultaneous deposit of different materials with a predetermined mixing ratio, resulting in a microstructure compositional shift that tailors the material properties at several regions of the produced component [1].

Japanese scientists originally proposed the notion of Functionally Graded Material (FGM) in the 1980s for high temperature uses in space aircraft [2]. Fabrication of FGM combines various materials in a single manufacturing operation to create metal-to-metal, polymer-to-metal, and metal-to-metal composite material combinations at specific places. There are different ways to produce FGMs using Additive Manufacturing (AM) techniques such as binder jetting, material jetting, powder bed fusion, etc. One of the well-known AM methods to fabricate FGM is the Fused Filament Fabrication (FFF) process. FFF techniques uses the layer- by-layer deposition method to build 3D objects. This method benefits from the wide variety of materials that are readily available. like thermosets, metal-filled thermoplastic filaments, composites, and flexible elastomers [3]. It is a low-cost technology and uses a low-cost material compared to other AM methods [4].

like thermosets, metal-filled thermoplastic filaments, composites, and flexible elastomers [3]. It is a low-cost technology and uses a low-cost material compared to other AM methods [4]. Conventional multi-material parts are fabricated using a separate extrusion nozzle for each material in FFF technology, limiting the smooth transition of materials. Swapping materials without altering the extrusion head is a benefit of the FGAM method over traditional AM. The procedure of depositing numerous materials using the FFF approach is shown in Fig. 1.240Fig 1: Description of the FFF process used in this study [12]A limited number of studies were performed to design, manufacture, and characterize the behavior of FFF-made FGMs. Doubrovski et al. [5] demonstrated a new work path, which began with defining and representing part geometry and material attributes. Using the bitmap approach, Material attributes are converted to the material composition.

Using the bitmap approach, Material attributes are converted to the material composition. They displayed workflow in a method that allows designers to improve the performance and usefulness of their products. Ituarte et al. [6] presented research on the multi-material binder jetting technique for designing and producing functionally graded structures. They created and evaluated digital material tensile specimens using a variety of processing conditions.the mechanical characterization of digital materials such as Young’s modulus, tensile strength, and elongation at break. To describe and simulate FGM structures, prescribed and automatic methodologies were used. Interface qualities of additively built multi-material components have been studied in several studies. Freund et al. [7] FFF technique explored the parameters influencing the interface strength of additively built multi-material components.

Freund et al. [7] FFF technique explored the parameters influencing the interface strength of additively built multi-material components. In this study, roll peeling tests and peeling resistance of numerous stiff and flexible materials were carried out to identify and quantify significant adhesion mechanisms. The findings indicated that materials’ polarity and mechanical interlocking significantly impacted the interface strength of components made using the FFF process. Another one-of-a-kind study was conducted by Garland et al. [8] outlined the difficulties and solutions experienced when developing and producing FGM parts using a commercially available FFF 3D printer. Results of the study contributed to the fabrication of multi-material structures with the optimal material distribution. Guessasma et al. [9] studied the manufacturing of multi-material objects using droplet-based AM methods.

Guessasma et al. [9] studied the manufacturing of multi-material objects using droplet-based AM methods. A two-extrusion head printer was used to create a composite material made of TPU and acrylonitrile butadiene styrene (ABS). 3D imaging methods were used to accomplish the porosity investigation. It was found that the interface quality and orientation influenced multi-material printed parts.The authors employed a regression-based material model forAnother one-of-a-kind investigation was carried out by Lopes et al., [10] which examines the mechanical properties of multi-material printed parts, with particular attention to the interface zone's geometrical boundaries. Various tensile specimens were evaluated using a mixture of PLA, TPU, and PET materials. The results revealed that multiple materials in pairs of PLA-TPU and PLA-PET exhibit a more significant reduction in Young's modulus and tensile strength.

The results revealed that multiple materials in pairs of PLA-TPU and PLA-PET exhibit a more significant reduction in Young's modulus and tensile strength. In this case, in addition to the physical issue of the boundary interface, there is also the chemical issue to address. There is a limited number of studies on the functionally gradient of AM materials; and the impact of processing parameters are layer height, printing speed and plane of orientation on the tensile and 241fatigue strength of AM plastic components [13-15]. This research study investigates the tensile and fatigue characteristics of two polymers having gradient compositions. Furthermore, the interface bonding strength of tensile specimens is investigated by mixing two materials at various concentrations. For fatigue analysis, the obtained UTS is utilized to set stress levels. At 80% of nominal UTS of each specimen type, tensile–tensile fatigue experiments with a stress ratio of R = 0.1 were performed.

At 80% of nominal UTS of each specimen type, tensile–tensile fatigue experiments with a stress ratio of R = 0.1 were performed. For fatigue testing, different printing planes are explored.Materials and MethodsMaterial SelectionTo examine amorphous polymers with various material characteristics, several materials were used. Amorphous polymers such as Thermoplastic Polyurethane (TPU) and chopped carbon fiber Polyethylene Terephthalate Glycol (CF-PETG) were employed in this study, Due to lack of research study on this material, it is being used. This type of FGM material can be applied to robotic grippers and prosthetic socket applications.

This type of FGM material can be applied to robotic grippers and prosthetic socket applications. Parameters used for printing the specimens as shown in Table1.Table 1: Fixed processing parametersParameters Nozzle temperature (oC) Bed temperature (oC) Infill density (%) Infill pattern Layer width (mm) Layer height (mm) Printing speed (mm/sec)Values 245 75 100 Line (0/90) 0.35 0.23 15TPU is a highly flexible material with good tensile strength. It is very resistant to abrasion, oil, and grease. The appropriate parameters for printing TPU specimens are bed temperature of 60 oC, nozzle temperature of 220 oC, and printing speed should below 15 mm/s. TPU Material occasionally won’t mix correctly with CF PETG if the printing speed is greater than 15mm/s. The most critical aspect is that the retraction settings be turned off.

TPU Material occasionally won’t mix correctly with CF PETG if the printing speed is greater than 15mm/s. The most critical aspect is that the retraction settings be turned off. It is utilized in several applications since it combines the advantages of both plastic and rubber, such as aerospace, automotive, wires, mobile covers, cables, etc. [16-19]. CF-PETG has high impact strength. It has good chemical and impact resistance properties. It is utilized in various applications, including prototypes and models, food and drink containers, medical and pharmaceutical applications, etc. Some materials are commonly used in both regular life and the 3D printing industry.

Some materials are commonly used in both regular life and the 3D printing industry. Table 2 shows the typical characteristics of TPU and CF- PETG polymers.Table 2: Typical properties of TPU and CF-PETGPropertiesValuesMelting Point (◦C) Density (g/cm3) Tensile strength (MPa) Glass transition temperature (◦C) Bed temperature(◦C)CF-PETG 230-250 1.23 45.8 82 70-75TPU 210-230 1.23 26 60 60-65By managing the deposition feature of materials at the voxel scale, voxel-based three- dimensional printing allows to producing digital structures at the mesoscale [11]. There is a variety 242of ways to create the FGM specimen. FGM specimens were designed using a voxel. A voxel is a graphic data unit that defines a point in three-dimensional (3D) space.

There is a variety 242of ways to create the FGM specimen. FGM specimens were designed using a voxel. A voxel is a graphic data unit that defines a point in three-dimensional (3D) space. Pixels create a point in two- dimensional (2D) space, initially, the object must be made in a CAD model, prepped, and exported into the Standard Tessellation Language (STL) file format, and then uploaded into slicer software, which generates G-code files that 3D printers can read. In STL format, the CAD tool produces the geometrical information of the model. Voxelizer slicer imports an STL file. The design of functionally graded components utilizing a voxlizer-based technique may handle heterogeneous material information. The Voxlizer slicer allows the user to alter the distribution of materials based on their volume.

The Voxlizer slicer allows the user to alter the distribution of materials based on their volume. In this research study a Zmorph desktop FFF-3D printer was used to fabricate FGM components with a gradient interface, as shown in figure 2.Fig 2: Gradient transition of FGM interfaces, (a) conventional versus FGM material combination, (b) fabricated components with different material combinations.Specimen fabrication and preparationThe corrosion of metals and alloys standard ISO 11782-1 was applied for these specimens, as shown in Fig.3. The geometry is Euro standard version specifically for the fiber-reinforced plastics. The length is 160 mm, with a 5 mm thickness and a 20 mm breadth.

The geometry is Euro standard version specifically for the fiber-reinforced plastics. The length is 160 mm, with a 5 mm thickness and a 20 mm breadth. At Gauge length, the curvature radius is 100 mm.Fig 3: Dimension of tensile and fatigue test specimen as per ISO 11782-1243Experimental SetupTensile and fatigue tests were carried out on a Testresources 810E4 load frame with a 15 KN load cell in line with ISO 11782-1 standard test procedure for uniaxial fatigue characteristics of plastics in this investigation. The specified stress level was maintained using a PID controller. The strength of FGM samples is comparable to CF PETG and TPU materials, as seen in Fig. 4. The test can be performed at a frequency of up to 25 Hz, still not more than 5 Hz, according to the standard.

4. The test can be performed at a frequency of up to 25 Hz, still not more than 5 Hz, according to the standard. Lower frequencies, on the other hand, were chosen to prevent the hydraulic actuator from being overheated.Fig 4: Experimental setup for fatigue and tensile testResults and DiscussionAnalysis of Tensile behaviorExperimental characterization of tensile and fatigue test samples has been carried out to investigate the behaviour of FGM composites under various structural loads. The tensile test results, in the form of stress–strain diagram, are given in Figure 5 and in Table 3. Engineering stress-strain diagrams for CF-PETG, TPU, and FGM printed specimens.Table 3: Ultimate Tensile strength (MPa) of 3D printed specimensSpecimen No. 1 2 3 AverageCF- PETG 40.43 42.53 43.29 42.08FGM 14 14.01 14.85 14.28TPU 8.6 8.74 9.05 8.79Table 3 shows the tensile strength of the FGM is more than TPU but less than the CF-PETG.

1 2 3 AverageCF- PETG 40.43 42.53 43.29 42.08FGM 14 14.01 14.85 14.28TPU 8.6 8.74 9.05 8.79Table 3 shows the tensile strength of the FGM is more than TPU but less than the CF-PETG. The ultimate tensile strength for the FGM specimen is around 14 MPa. The characteristics of CF- PETG and TPU were passed down to FGM specimens. FGM is one of the methods for improving the strength of FFF Specimens' weak bonding layer.244454035CF PETG FGM TPU) a P M3025(s s e r t S201510500%5%10%15%Strain (%)Fig 5: Stress-strain Graph of CF/PETG, FGM and TPUTensile test results in Figure 6 revealed the variation of material properties from neat TPU to CF/PETG. The graph shows the decreasing trend in material’s strength and stiffness properties from neat TPU to CF/PETG fabricated on the XY plane.

The graph shows the decreasing trend in material’s strength and stiffness properties from neat TPU to CF/PETG fabricated on the XY plane. This indicates that the mechanical behavior of the structure can be tunable in FFF made FGM parts to obtain desired properties at desired locations.4035CF_PETG - 100%3080P_20T) a P M2560P_40T40P_60T(s s e r t S201520P_80TTPU - 100%10500%2%4% Strain (%)6%Fig 6: Stress and strain graph of various concentrations of CF-PETG and TPU materials245Fatigue test resultsThere is a scarcity of research articles on fatigue study of FGM and FFF specimens, as stated in the introductory portion of the paper. The tension-tension fatigue test was performed at 80% of the specimen’s maximum load with a stress ratio of 0.1 and a frequency of 3 Hz. To understand the effect of orientation process parameters, two different plane orientations (XY-0) and (XY-90) were tested.

To understand the effect of orientation process parameters, two different plane orientations (XY-0) and (XY-90) were tested. Figure 7 shows the number of fatigue cycles for the various raster orientations of FGM specimens. Samples printed on XY-0 directions shown a higher fatigue life cycle than the XY-90. This is because 3D printed parts are strong if a load is applied along with the printing orientation.600500s e l c y c400f o r e b m u N3002001000XY-0 CyclesXY-90 CyclesFig.7: Fatigue cycles of the specimens fabricated on flat (XY-0) and (XY-90) orientationsIn this research study, the analysis of variance (ANOVA) method was used to understand the effect of the processing parameters on the fatigue behavior of the FGM samples. The ANOVA table is given below it shows that the orientation has a statistical significance with a P-value of 0.03 which is less than 0.05.

The ANOVA table is given below it shows that the orientation has a statistical significance with a P-value of 0.03 which is less than 0.05. This indicates that the change in printing orientation will statistically impact the fatigue life of functionally graded samples.Table 5: ANOVA TableSource Model Error Corrected Total Source Orientation SpecimenDF Sum of Squares Mean Square 4 3 7 DF 1 352915 5887 58802 Type I S S 29282 2363313228.75 1962.3333Mean Square 29282 7877.66667F-Value 6.74F-value 14.92 4.01Pr > F 0.0744Pr > F 0.0307 0.1418The results in Table 5 match the results of figure 7, showing the effect of the printing plane orientation. Accordingly, the XY-0 raster orientation specimen had the most extended fatigue life at the 80 percent UTS stress level, with 477 cycles to failure.

Accordingly, the XY-0 raster orientation specimen had the most extended fatigue life at the 80 percent UTS stress level, with 477 cycles to failure. Where DF, SS and Pr are Degree of freedom, Sum of Square, and Probability.246MicrostructureScanning electron microscope (SEM) analysis was performed to examine fracture mechanisms and microstructural morphology of the composite specimens. The result of the SEM analysis revealed that fiber orientation and distribution influenced the fracture behavior of the FGM samples, as shown in Figure 8.Fig. 8: Dog bone sample of FGM with 10% of short carbon fiber before fatigue testing(A) Short carbon fiber (SCF) (B) cross viewFiber breaking, debonding, and fiber pull-out are the fracture modes following fatigue testing. The cyclic load causes fiber breakage and debonding. Some fibers are damaged, while others are yanked out of the matrix. In Figure 9, broken fibers are seen surrounding the fractured area.Fig.

The cyclic load causes fiber breakage and debonding. Some fibers are damaged, while others are yanked out of the matrix. In Figure 9, broken fibers are seen surrounding the fractured area.Fig. 9: FGM samples after fatigue testing (A) cross-section (B) side view247ConclusionThis study used experimental testing, SEM analysis, and the ANOVA method to evaluate the mechanical characterization, tensile and fatigue behaviors of FFF-made FGMs.

The current paper summarizes the findings as follows:The FFF process successfully fabricated the FGM specimens using the single nozzle multi-material printing technology.It was found that the mechanical behaviour of TPU-CF/PETG composites can be tuneable by mixing the specific ratio of each material.The XY-0 printing plane orientation of the FGM specimen had the most extended fatigue life as compared to the XY-90 direction.The study also revealed that FGM could be used to improve the interlaminar shear strength of weak FFF parts.Microstructural analysis indicated that the primary failure mechanism was responsible for a weak fiber-matrix interface when applying the cyclic fatigue loading.Future WorkThe scope of the future studies is given below:Improving the software capabilities to design highly complex FGM patterns is needed.Understanding the fracture behaviour of multi material interfaces under bending fatigue.Fatigue modelling and characterization.AcknowledgmentThe authors greatly acknowledge the support provided by the Center for Manufacturing Research.References[1]Loh, G.H.; Pei, E.

3D Print Addit Manuf 2016; 3:71–8. https://doi.org/10.1089/3dp.2016.0026[12] Hasanov, Seymur & Gupta, Ankit & Nasirov, Aslan & Fidan, Ismail. (2020). Mechanical characterization of functionally graded materials produced by the fused filament fabrication 923-935. 10.1016/j.jmapro.2020.09.011.process.Journalof Manufacturing Processes.58.[13] Hasanov S, Gupta A, Alifui-Segbaya F, Fidan I, Hierarchical homogenization, and experimental evaluation of functionally graded materials manufactured by the fused filament fabrication process. Composite Structures 275 (2021): 114488.[14] Hasanov S, Gupta A, Nasirov A, Fidan I, Mechanical characterization of functionally graded materials produced by the fused filament fabrication process. J Manuf Process 2020; 58:923–35.[15] Nasirov A, Hasanov S, Fidan I, Technology E. Prediction of Mechanical Properties of Fused Deposition Modeling Made Parts using Multiscale Modeling and Classical Laminate Theory. Proc. 30th Annu. Int. Solid Free.

PERIMENTAL ANALYSIS ON AN ADDITIVELY MANUFACTURED ABS LIVING HINGECassandra Gribbins*, Heidi M. Steinhauer, Ph. D*†Department of Mechanical Engineering, Embry-Riddle Aeronautical University, Daytona Beach, FL 32114 †Department of Engineering Fundamentals, Embry-Riddle Aeronautical University, Daytona Beach, FL 32114 AbstractA study on the plastic behavior of an additively manufactured acrylonitrile butadiene styrene (ABS) living hinge was conducted using a MakerBot 2X. Initial research included numerical and analytical linear analyses on a typical living hinge design. This paper introduces the portion of the research that explores the application of traditional design practices to entry-level additive manufacturing machines. Tensile testing for material properties was conducted to refine the numerical model. Experimental rotational testing was conducted for data on the non-linear, plastic behavior experienced during application.

Experimental rotational testing was conducted for data on the non-linear, plastic behavior experienced during application. Verification of the numerical model with experimental results will be used to guide future work on exploring alternate design geometries that leverage the advantages of additive manufacturing’s design freedom for smoother stress distribution on the hinge.IntroductionInducing flexural capabilities within a single plastic piece is often executed through the utilization of a living hinge design. Bending is achieved by creating a relatively thin section of plastic between two larger, rigid regions. Incorporating living hinges in a design reduces part count which can result in lower overall costs and assembly time [1].Initial research compared numerical and analytical analyses of a traditionally design living hinge against an alternate design.

There was a large percent difference between stresses from the two solutions due to the analytical solution being conservative and the numerical solution overestimating the non-linear results. The initial research also indicated the hinges behaving in the plastic region under a small deformation of 10°. Comparing the calculated stresses with the yield strength to determine hinge behavior showed that all cases acted within the plastic region under the enforced deformation.Further research suggestions to refine the results included testing for more applicable material properties for analysis and failure criteria of the living hinge. The initial research adjusted the yield strength of bulk ABS with respect to experimental investigations from Ahn et al. [2] and Rodriguez et al. [3] that demonstrated fused deposition modeled (FDM) ABS having 65 to 75% yield strength of injection molded ABS, resulting in a usable yield strength of 28.6 MPa.

[2] and Rodriguez et al. [3] that demonstrated fused deposition modeled (FDM) ABS having 65 to 75% yield strength of injection molded ABS, resulting in a usable yield strength of 28.6 MPa. In Ahn et al.’s study [2] the common dogbone-shaped sample defined by the ASTM D638 standard was prone to break at the radii due to stress concentrations induced by gaps in the toolpath generation. The ASTM D3039 standard was then used for tensile testing [2]. Alternately, Lee and Huang [4]conducted fatigue testing using the ASTM D638 standard and did not report any interference of results when a few samples fractured at the radii.In Rodriguez et al.’s [3] research, the ASTM D3039 standard was also used to conduct tensile testing. They concluded a more significant reduction of 22 to 57% in strength relative to ABS monofilament [3]. This reduction is due in part by voids formed during the process. Default building parameters inherently resulted in voids within generated toolpath previews.

This reduction is due in part by voids formed during the process. Default building parameters inherently resulted in voids within generated toolpath previews. Hossain et al. [5] demonstrated a visual feedback method of adjusting building parameters based on a magnified optical image of the printed part as modifications of parameters using the toolpath preview resulted in gaps that were not identified within the preview.Toolpaths and other building parameters like build orientation affect the strength of the part inducing an anisotropic nature in FDM created parts. Properties can be considered isotropic within the x-y plane, while strength in the z-direction is measurably less due to the tendency to delaminate between layers [6].Gibson et al. [6] investigated a traditional design of living hinge using a PolyJet 3D printing technology that utilizes photopolymer material [6]. Initial results indicated success but encourages further testing for heavy use.

Initial results indicated success but encourages further testing for heavy use. Stratasys, Ltd [7] reports manufacturing an FDM living hinge that lasts up to thousands of cycles. Build recommendations include printing living hinges in a vertical build orientation for the best hinge durability as shown in Figure 1. AM living hinges still have room for improvement as traditionally injection molded polypropylene living hinges that can last millions of cycles [8].Figure 1 FDM living hinge printed in the vertical build orientation [7]ExperimentsMaterial TestingTensile testing for material properties was conducted to refine the material properties used in the CATIA V5 numerical model. Young’s modulus influences how the stress is determined from the deformation/strain on the part.

Young’s modulus influences how the stress is determined from the deformation/strain on the part. The yield strength sets the failure limit for designing a living hinge to act within the elastic region.Tensile testing was conducted using the Tinius Olsen Model 290 Lo-Cap Universal Testing Machine with a 133,500 N load capacity. A Tinius Olsen S-400-2A extensometer was used to obtain strain data during tensile testing as shown in the testing setup in Figure 2. The strain rate applied was variable with an average of 20 mm/min.Initial tensile testing was performed to determine proper design geometry between ASTM D3039 [9] and ASTM D638 Type I [10]. The specimen adhering to ASTM D3039 fractured within the grips while the ASTM D638 specimen fractured at the base of the radii similar to the results reported by Ahn et al.

The specimen adhering to ASTM D3039 fractured within the grips while the ASTM D638 specimen fractured at the base of the radii similar to the results reported by Ahn et al. [2].ExtensometerFigure 2 Tensile testing setup in the Tinius Olsen with extensometer attachedFor the second iteration, 1/8” thick aluminum tabs were applied to the ends of the specimen for better grip and to prevent fracturing within the grips. Another tensile test resulted in the ASTM D3039 specimen fracturing at the location of the tabs. The ASTM D638 specimen again fractured at the base of the radii. Figure 3 shows the second iteration specimen failure.Figure 3 Second iteration tensile testing specimen showing ASTM D639 at top fractured at the radii and ASTM D3039fractured at the tabThe ASTM D638 design geometry was chosen for further testing as it did not fracture within the tab.

The crazing displayed along the narrow length of the specimen was more evenly distributed for the ASTM D638 sample, as shown in Figure 4a. The crazing in the ASTM D3039 specimen was more concentrated toward the location of fracture as shown in Figure 4b.Figure 4 Close-up image of crazing in a) ASTM D638 and b) ASTM D3039Further refinement of the testing specimen included elongating the ends for more grip and decreasing the thickness to compensate for the additional thickness provided by the tabs. The testing area remained the same otherwise. Figure 5 displays the quarter base sketch used to generate the tensile specimen. The overall thickness of the part was 3.5 mm.Figure 5 CATIA V5 base sketch of one quarter of the tensile specimen (units in mm)The CATIA V5 part was exported to an (stereolithography) STL file with a sag size of 0.001 mm and imported into MakerBot Desktop to generate the toolpath data for printing.

The specimen were all arranged to build in the vertical orientation as shown in Figure 6. The specimen were created in the vertical build orientation as the living hinges were also printed vertically. The same print orientation as the living hinge would provide material properties that represent the hinge structure.Figure 6 MakerBot Desktop Home View position of tensile specimen in vertical print orientation with coordinate system shown under the Change Position box (support structure not shown)As a result of the vertical orientation, support material was generated to support the part underneath the curve. There is limited control over generation of toolpath direction under the main options so techniques on optimizing building parameters were not utilized. The standard print profile with default values from MakerBot Desktop were used although the infill was changed to 100% for a solid part.

The standard print profile with default values from MakerBot Desktop were used although the infill was changed to 100% for a solid part. Toolpath preview was reviewed and discovered that MakerBot Desktop automatically generates a 45°/-45° alternating toolpath for the outer three layers on a part along the X-Y plane and switches to 0°/90° toolpath for the layers in between. Figure 7 illustrates a layer of the tensile specimen that combines the 45°/-45° toolpath for the outer layer of the narrow length of the part and the continuation of the 0°/90° for the left wide tab end.

The standard setting also resulted in the presence of voids within the structure as indicated by the print preview.Figure 7 MakerBot Desktop top view of the print preview illustrating 0°/90° toolpath on the left and 45°/-45° for the outer layerof the narrow length on the rightThe original cross-sectional area to be used in material property calculations was obtained by taking the average of the width and thickness measurements from the narrow section of the tensile specimen. The measurement locations are shown by the black markings in Figure 8 using Pittsburg 6” digital calipers with a resolution of 0.01mm.Figure 8 Third iteration tensile testing specimenApplication TestingExperimental testing was planned to be conducted for data on the non-linear, plastic behavior experienced during application, but the micro-tensile machine to be used was unavailable.

Fabricating of a living hinge was carried out to determine machine capabilities.The model containing the traditional living hinge design, as shown in Figure 9, was created in CATIA V5. The minimum thickness of 0.3 mm for the hinge thickness would not be rendered by the MakerBot Desktop software in the print preview window when oriented in the vertical printorientation. The hinge thickness was increased by 0.1 mm increments until the MakerBot Desktop software rendered the hinge section in the print preview.Figure 9 Traditional living hinge design [1]The minimum hinge thickness that MakerBot Desktop would render was 0.6 mm, indicating geometry less than 0.6 mm cannot be printed. The hinge length was also adjusted to 3 mm as demonstrated by the living hinge experimentation by Goenka [11]. The lower recess of the hinge was modified from a semi-circle shape as illustrated in Figure 9 to a straight lower fiber with a 0.2 mm radius.

The lower recess of the hinge was modified from a semi-circle shape as illustrated in Figure 9 to a straight lower fiber with a 0.2 mm radius. The uniform hinge thickness was utilized to follow design suggestions by Stratasys, Ltd [7]. The final design for a printable living hinge on the MakerBot 2X is shown in Figure 10.Figure 10 CATIA sketch of half a living hingeThe solid model in CATIA was characterized as a material with the Young’s modulus and yield strength determined from tensile testing as shown in Figure 11. The remaining properties were obtained from CES EduPack 2013 [12]. Isotropic material was selected as the anisotropic option contained many necessary properties that were not available.Figure 11 Material properties used in the FEACATIA V5’s Generative Structural Analysis workbench was used to perform a finite element analysis (FEA) on the solid model.

The left face was fully constrained and an enforced displacement was applied on the right end surface as shown in Figure 12. A vertical displacement of 4.5 mm was applied for a comparable 10° rotation. The enforced displacement was defined to use the axis system at the center of the hinge to encourage circular bendingFigure 12 FEA case modelAn overall part mesh of 0.4mm was generated by CATIA based on the part dimensions. A local mesh size around the hinge area was refined until the resultant maximum von Mises stress was with within 3% of the previous 3 cases. Figure 13 displays the local mesh refinement of 0.16 mm about the hinge.Figure 13 FEA local mesh refinementFurthermore, five living hinge samples were printed and measured for comparison between the theoretical dimensions and the resulting print after shrinkage, which for ABS is generally about 2% [13]. All hinge specimen were printed in the vertical orientation as recommended by Stratasys, Ltd [7].

All hinge specimen were printed in the vertical orientation as recommended by Stratasys, Ltd [7]. The print settings were set to the low/fast setting for MakerBot adjusting only the infill to 100%, the number of shells to 1, and reducing the layer height to the standard setting of 0.2 mm.Results and DiscussionMaterial Testing ResultsThe fracture surface of the ASTM D3039 specimen from the second iteration was examined using a FEI Quanta 650 Scanning Electron Microscope (SEM) under the low vacuum setting. Voids similar to the one shown in Figure 14 were discovered. The depth and smooth walls of the void suggests that it was created during manufacturing and not a microvoid from part of the crazing.Figure 14 Scanning Electron Microscope (SEM) image of tensile testing specimen fracture surface illustrating voidmeasurement of 47.28 µm by 30.47 µmThe measurements for the third iteration of tensile specimen are shown in Table 1.

The results did not show any pattern on how the values differed at various areas across the narrow section.

The thickness was considerably smaller than the theoretical and is attributed to the tendency for ABS to shrink approximately 2%.Table 1 Tensile testing measurements for determination of usable cross-sectional areaBuild OrientationSpecimen NumberW1W2W3W4W5Overall Average (mm)SDT1T2T3T4T5Overall Average (mm)SDCross-sectional Area, A0 (mm2)45.50I13.0212.9712.9612.9912.9812.980.02063.363.373.373.363.363.360.004943.68II13.1613.1913.1913.1713.1213.170.02583.473.493.513.523.523.500.019446.11III13.1713.1513.1213.0913.1813.140.03313.373.373.363.363.363.360.004944.21IV13.0213.0613.0412.9813.0013.020.02833.413.413.413.423.413.410.004044.42V13.0813.0713.0813.0613.0713.070.00753.413.413.413.413.413.410.000044.58Sectional Average13.0913.0913.0813.0613.0713.083.403.413.413.413.413.4144.60SD0.07280.08560.08610.07790.08310.04340.04900.05930.06540.0653VerticalWidth (mm)Thickness (mm)Theoretical13.003.50From the third iteration, all five specimen fractured at the yield strength displaying brittle behavior.

The failed specimen are shown in Figure 15. As shown, the fracture occurred at the base of the radii but still across an area similar to that measured across the narrow length. A lower strain rate may encourage more plastic behavior.Figure 15 Tensile test specimen failureThe engineering stress vs strain curves from the third iteration are displayed in Figure 16. The curve displays the same characteristics as Rodriguez et al. [3] for a test specimen with fibers running perpendicular to the load direction. The fracture toughness of the material may have an influence on the brittle fracture of the specimen if the microvoid is greater than the allowable flaw size. The sample V3 appears to have experienced some slippage in the extensometer near the yield point of the curve, otherwise the results across the five specimen were consistent. The proportional limit of the curves were all around 15 MPa.

The proportional limit of the curves were all around 15 MPa. The Young’s modulus was calculated using points at the beginning and end of the modulus line as recommended by an instruction pamphlet from Tinius Olsen [14].Engineering stress vs strain40.00) a P M35.0030.00(s s e r t S g n i r e e n i g n E25.0020.0015.0010.005.000.000.00000.00500.01000.0150 0.0200 Strain (mm/mm)0.02500.03000.0350Figure 16 Engineering stress vs strain curves from the tensile specimenThe summary of determined material properties for each specimen and the overall average is shown in Table 2. The Young’s modulus average of 2141 MPa and tensile strength average of 35.16 was used in the adjusted material properties for the CATIA V5 model.V1V2V3V4V5CES EduPack 2013 [12] cites a range of values for each material property and for injection molded ABS the Young’s modulus is stated as ranging from 2210 to 2620 MPa.

The determined average for FDM ABS of 2141 MPa is 97% of the lower end of the injection molded material. The injection molded range for yield strength is 42 to 46 MPa resulting in the experimental yield strength of 35.2 being 84% of the lower end. The yield strength is a great improvement from previous material studies and can be attributed to the improvement of toolpath generating programs and overlapping fibers.Table 2 Summary of material properties obtained from tensile testingSpecimenYoung's Modulus (MPa)Yield Strength (MPa) Yield Strain (mm/mm)I1983.5936.560.0321II2219.8234.010.0229III1985.1634.930.0297IV2146.1136.030.0271V2370.6734.270.0243Average2141.0735.160.0272SD147.010.990.0034Application Testing ResultsFrom the results of the tensile testing, the failure criteria is defined by stress beyond the yield strength of 35.2 MPa.

The translational displacement diagram shown in Figure 17 verifies that the hinge deformed as expected about the center of the hinge.Figure 17 Translational displacement vector from CATIA V5 Generative Structural Analysis workbenchFigure 18 shows the von Mises stress distribution across the hinge with the yield stress of 35.2 MPa set as the maximum limit. High stress is experienced across the entire hinge with the highest stress experienced is 37.7 MPa under the lower portion of the hinge. This is above the yield stress under a 10° deformation. From the material testing, stress beyond the yield would result in fracture. This current design for a living hinge would not be practical for use with such a minimal operating range.Figure 18 von Mises stress distribution from CATIA V5 Generative Structural Analysis workbenchFigure 19 displays the printed living hinges with the rearmost hinge placed in the vertical build orientation.

It is interesting to note that due to the small design geometry, the lower portion of the hinge resulted in a curve close to the traditional design as shown in Figure 9.Figure 19 Manufactured living hinges with the rearmost hinge shown in the vertical build orientationThe results of the manufactured living hinge sample measurements are shown in Table 3. The length and width of the hinge dimensions were all lower than the theoretical while all of the measured thicknesses were all above the theoretical.

The same pattern appeared in the overall dimensions with length and width both being lower than theoretical while thickness measured either at or slightly above the theoretical.Table 3 Measurements for a set of traditional designed living hingesLength (mm)Width (mm)Thickness (mm)Length (mm)Width (mm)Thickness (mm)3.007.400.6050.807.402.30I2.757.090.9450.537.332.31II2.717.110.9350.467.272.30III2.757.130.9450.537.352.31IV2.727.110.8550.467.272.30V2.777.120.9250.487.252.30Average2.747.110.9250.497.292.30SD0.02190.01330.03380.03190.03880.0049Hinge DimensionsOverall DimensionsTraditionalTheoreticalSpecimenWhen taking into account the expectation of ABS shrinking approximately 2%, the overall length and width averages are reasonable measurements while the length and width for the hinge portion are greater than the expected 2% shrinkage minimum.

The effects of small hinge dimensions, as well as its location in the center of the part, may have contributed to a greater than expected shrinkage.The theoretical value for hinge thickness, while shown to be large enough to manufacture in the print preview, was too small for the printer to create resulting in the apparent minimum thickness that can be printed as approximately 0.9 mm, although printers of similar capabilities can print to smaller values. The overall thickness resulted in nearly theoretical values.Conclusion and Future WorkTensile testing showed increased material property values from previously approximated tensile strength, 28.6 MPa to 35.2 MPa. The characteristics of the stress-strain curve displayed brittle behavior that can be attributed to the possibility of microvoids affecting the fracture toughness of the material. Future work into testing and determination of the fracture toughness property for FDM ABS is suggested.

Future work into testing and determination of the fracture toughness property for FDM ABS is suggested. As advances in toolpath generation further minimize voids, determining the allowable flaw size is important when designing with respect to the critical stress of the structure.As recommendations for future work, additional research for improving accuracy of small structures like living hinges using a MakerBot 2X would involve exploring the advanced options of MakerBot Desktop. Adjustment of the advanced options requires understanding of MakerBot terminology of the different parameters and how they affect the overall build.Alternate designs for living hinges would be a study of interest for further research. The traditional design for a living hinge allows too high of a stress concentration in the small area as indicated by the FEA.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperExperimental and Computational Study of Area Printing™ Additive Manufacturing: Inconel 718 and M300 Maraging Steel Density ImprovementB. Fotovvati, S. Shrestha, N. Ferreri, and N. DuanmuSeurat Technologies Inc., Wilmington, MA, USA 01887AbstractThe low manufacturing speed of laser-powder bed fusion (LPBF) additive manufacturing has hindered its adoption in conventional manufacturing methods. Large-area pulsed laser powder bed fusion (LAPBF), also known as "Area Printing™", has addressed this limitation by replacing the point laser with a large-area pulsed laser.

Large-area pulsed laser powder bed fusion (LAPBF), also known as "Area Printing™", has addressed this limitation by replacing the point laser with a large-area pulsed laser. Each pulse melts a region of the powder bed in the order of square millimeters, which enables scalability without the sacrifice of resolution and high throughput at an equivalent or better quality (in part due to a lack of spatter) when compared to conventional LPBF methods. In this study, process parameters are optimized to achieve near-full- density parts, and a computational model is developed to understand the multi-physics governing the process. It is observed that the shallow depth and high aspect ratio of the melt pool lead to a unidirectional solidification front extending along the build direction where grains grow epitaxially, and highly directional microstructures are created.Keywords: Selective Laser Melting; Area Printing™; Thermo-fluid Modeling; Inconel 718; M300 Maraging Steel10801.

IntroductionLaser-based powder bed fusion (LPBF), also known as selective laser melting (SLM), or direct metal laser sintering (DMLS), is one of the most popular types of metal additive manufacturing (AM). LPBF employs metal powders selectively melted in a layer-wise manner using a high-intensity laser beam in correspondence with a sliced, computer-aided design (CAD) model. LPBF is widely used in various industries, including medical, automotive, and aerospace. Some examples of these applications include patient-specific metal implants [1], turbine blades [2], engine manifolds [3], and lattice structures [4]. The popularity of this manufacturing technique in such industries is due to the many advantages that this method offers: design freedom, reduced manufacturing time, minimized material waste, sustainability benefits, etc. [5].

[5]. However, LPBF's low production rate coupled with the typical capital and operational costs, has limited its adoption over conventional manufacturing methods. To address this challenge, Seurat Technologies has developed large-area pulsed laser powder bed fusion (LAPBF), also known as "Area Printing™," which replaces the point laser in conventional LPBF with a large-area pulsed laser. LAPBF melts on the scale of several millimeters in a single pulse with a repetition speed of up to 40 Hz. This can result in significantly shorter manufacturing times and lower final production costs. The large- area laser beam also minimizes spattering issues commonly encountered in the conventional LPBF process, as there is no deep turbulent melt pool to create strong recoil pressures and eject particles.Several studies have investigated the effect of laser parameters, such as laser power and scan speed, on the quality and properties of LPBF-fabricated parts.

Research studies show changing laser power and scanning speed have different effects on pore formation in IN718 [6,7]. Additionally, Wang et al. [8] optimized the LPBF process parameters to get high-density IN718 samples and investigated the effects of heat treatments on microstructure and mechanical properties. Using a laser power of 150 W, a scanning speed of 550 mm/s, and a 45 % overlap rate (yielding an energy density of 158 J/mm3), they achieved a relative density of 99.8%. Liu et al. [9] explored the effects of the LPBF process parameters on the microstructure of IN718 samples and observed that as the energy density decreases, the aspect ratio of the columnar grains becomes lower due to insufficient epitaxial growth. They concluded that refined grains appeared due to the relatively fast solidification occurring in the relatively shallow melt pool. Georgilas et al.

They concluded that refined grains appeared due to the relatively fast solidification occurring in the relatively shallow melt pool. Georgilas et al. [10] investigated the effects of laser power and exposure time on the density of IN718 samples fabricated by pulsed LPBF. They achieved a maximum relative density of 99.8% with laser power of 200 W and exposure time of 110 μs. The grain structure of the as-built pulsed LPBF IN718 material was less directional (morphologically and crystallographically) than continuous LPBF.There is little work in the literature on the additive manufacturing of maraging steels. M300 and 18Ni300 are among the most popular maraging steels in the AM industry. Due to their excellent combination of strength and fracture toughness, maraging steels are mainly used in the tool and aerospace industries. Their very low carbon content gives them excellent weldability and high toughness [11]. Casalino et al.

Their very low carbon content gives them excellent weldability and high toughness [11]. Casalino et al. [12] showed that relative densities higher than 99.0% could be achieved for 18Ni300 maraging steel in the LPBF process by using laser powers higher than 90 W and a scanning speed lower than 220 mm/s. They also demonstrated that the hardness, the mechanical strength, and the surface roughness correlated positively to the part density. Krol et al. [13] investigated the effects of LPBF process parameters on the density of 18Ni300 maraging steel and achieved 99.3 % relative density with 200 W of laser power and a scanning speed of 340 mm/s.1081Mao et al. [14] used a relational model between the LPBF process parameters and the relative density of 18Ni300 maraging steel samples to obtain a relative density of 99.5% using laser power of 284 W and a scanning speed of 868 mm/s. Vishwakarma et al.

Vishwakarma et al. [15] studied the effect of build orientation on the microstructure and tensile behavior of LPBF-fabricated M300 maraging steel and achieved 99.6 % relative density using a laser energy density of 62.5 J/mm3. Fine grains with an average size of ~0.31 μm were observed in their as-built samples.In this study, the parameter optimization of the LAPBF process is presented to achieve near-full-density (>99.9 % relative density) parts using two popular metallic materials: Ni-based superalloy 718 (IN718) and M300 maraging steel. The aim is to obtain high-density parts. Additionally, a computational model is developed to better understand the underlying physics of the process, including thermal dynamics, fluid flow, and phase transformations, to provide insights into the relationship between process parameters and part quality.

Overall, this study aims to contribute to developing high-quality, cost-effective AM processes using LAPBF, which could have significant implications for various industrial sectors.2. Methodologies2.1. Experimental methodsIn Area Printing™, a laser containing millions of pixels is utilized to melt metal powder layers and fuse them to the area below it, manufacturing entire renderings at once in a single defined area. A set of diode lasers are used to preheat the powder bed before a single Nd:YAG pulsed laser performs the final melting of the area. Figure 1 illustrates a schematic of the pulsed laser beam from the source to the powder bed. Once the laser is created, it is shaped into a homogeneous square field called a "tile". A blue light projector then generates a projection of the area pattern the same size as the square laser beam, which is aligned with the laser beam.

A blue light projector then generates a projection of the area pattern the same size as the square laser beam, which is aligned with the laser beam. An optically addressable light valve (OALV) receives this light and horizontally polarizes the laser beam where blue and IR light pixels overlap and vertically polarizes them where they do not overlap. The vertically and horizontally polarized laser beams are then split, and the resulting area pattern is sent through a scanner system to the powder bed to melt the patterned field.IN718 and M300 maraging steel powders with particle size distributions of 15–37 µm and 15–45 µm, respectively, were used for the experiments. Each density-optimization print job contained a seven-by-seven grid of 10 mm x 10 mm x 10 mm density cubes on an Area Printing™ Prototype System. After print completion, the build plate was taken out, and the samples were cut from the build plate using a horizontal band saw.

After print completion, the build plate was taken out, and the samples were cut from the build plate using a horizontal band saw. The samples were then soaked in isopropanol in an ultrasonic cleaner for 30 minutes removal for density measurements. The density measurements are performed using the Archimedes method and an Ohaus Explorer™ Precision density kit. Micrographs were taken using an Olympus LEXT™ OLS5100 3D laser scanning microscope and processed using an in-house developed code for defect classification.1082Figure 1. Schematic of the pulsed laser beam used in the Area Printing™ process [16].2.2.Computational methodsThe integrated LPBF process simulation was implemented based on two steps: powder deposition using the discrete element method (DEM) and laser-powder interaction using computational fluid dynamics (CFD).

The powder deposition was modeled using the open-source DEM code LIGGGHTS (LAMMPS Improved for General Granular and Granular Heat Transfer Simulations) [17]. The material properties of IN718 used in the DEM simulation were density of 8.17 g/ml, modulus of elasticity of 20 MPa, Poisson's ratio of 0.29, sliding friction coefficients of 0.4, rolling friction coefficient of 0.1, and coefficient of restitution of 0.4.After the powder bed was obtained from the DEM simulation, it was saved as a Stereolithography (STL) file and imported to Flow-3D software version 11.2.6.4 (Flow Science, Inc., Santa Fe, New Mexico, USA). The governing equations of DEM and CFD are explained in our previous work [18,19]. Temperature-dependent material properties such as density, specific heat, thermal conductivity, and viscosity are used [20].Figure 2 shows the computational domain and dimensions.

Temperature-dependent material properties such as density, specific heat, thermal conductivity, and viscosity are used [20].Figure 2 shows the computational domain and dimensions. In this model, the powder particles are contactless or have minimal contact to increase the accuracy of slow powder bed heating with diodes. Hence, the mesh size varies from 1.2 µm to 5 µm in the powder region. Due to the lower mesh size, a laser illumination area of 0.3 mm × 0.3 mm was used to run the simulation in a reasonable time.1083Figure 2. Computational model of the powder bed.3. Results and DiscussionsThe results of the experiments and simulations are presented and discussed in the followingsub-sections.3.1. Experimental results 3.1.1. Effect of Laser Energy Density on Relative DensityThe observed effect of diode laser fluence on the porosity of area printed density cubes is presented in Figure 3.

Experimental results 3.1.1. Effect of Laser Energy Density on Relative DensityThe observed effect of diode laser fluence on the porosity of area printed density cubes is presented in Figure 3. The relationship amongst diode laser flux, pulse length, and porosity is illustrated in Figure 4 and constitutes a regression of the experimental data. The domain on each plot in Figure 3 and Figure 4 is divided into different regions based on the observed defect type. Notably, the near-full-density area for IN718 is significantly narrower than that of M300. The observed behavior for IN718 aligns with what is reported in the literature [21]. This is due to IN718 being prone to internal cracking at high energy densities [22]. No inflection point was observed in the porosity plot of M300 samples by increasing the applied energy density. More experiments at higher energy densities and comparing the fully dense samples at various energies are yet to be carried out.

Micrographs of representative M300 maraging steel samples from different regions of the porosity vs. energy density plot: (a) low energy density showing lack of fusion and insufficient melting defects and (b) high energy density resulting in near full density.Due to the nature of the Area Printing™ process, high-aspect-ratio melt pools are formed that are significantly wider and relatively shallower than the ones generally formed during the LPBF process. During the solidification process, cells grow epitaxially and perpendicular to melt pool boundaries [24]. As a result, in the case of having flat and horizontal melt pool boundaries for most of the part cross-section area, the favorable growth direction for most of the cells would be along the build direction.

The melt pool width of a full tile in the Area Printing™ process is more than 100 times larger than its depth, which makes the resultant microstructures of Area Printed™ parts highly directional along the build direction, as illustrated in Figure 7 for M300 maraging steel.1086BDFigure 7. A representative M300 maraging steel polished cross-section showing epitaxially grown, highly directional grains extending along the build direction and planar melt pool boundaries orthogonal to the build direction.3.1.2. Effect of Sample Size and Geometry on DensityIt is shown in the literature that the size and geometry of LPBF-manufactured parts may have an influence on the parts' properties [25]. To investigate this effect on the density of IN718 samples, cubes and cylinders were fabricated with different sizes but with the same laser parameters. The size of the samples ranges from 3 mm to 14 mm. Figure 8 shows these samples on a prototype machine’s build plate.

The size of the samples ranges from 3 mm to 14 mm. Figure 8 shows these samples on a prototype machine’s build plate. The laser parameters were chosen to have a high porosity to make the differences in density more pronounced. Figure 9 shows the porosity variations for samples with different shapes and geometries. The results suggest that smaller parts have significantly higher porosity; In other words, porosity is decreased by increasing the sample size. The trend then reaches a steady state for features bigger than ~2.5 mm. As the dominant cooling mechanism for samples is conduction through the solid material to the build platform, decreasing the sample size results in a decrease in the cooling rates. So, a reason for the increase in porosity1087by decreasing the sample size can be attributed to the accumulated heat in these samples due to slower cooling rates that can cause internal cracking in IN718 material, as discussed in section 3.1.1.

Therefore, in parts with complex geometries, different parameter sets can be used for regions with small features. Also, slightly higher porosity was observed for cubes of different sizes compared to cylinders, which can be attributed to cylinders having a more efficient cross section to dissipate the heat from the top layer to the build platform.Figure 8. IN718 samples with different sizes and geometries (the two top rows and the two bottom rows were used for this study).4.84.6Cubes Cylinders)%4.4( y t i s o r o P4.24.03.83.60.11.13.1 2.1 Sample volume (cm3)4.1Figure 9. Effect of sample size on porosity in IN718.10883.1.3. Effect of In-situ Laser Shock Peening and Tile Overlap on DensityAccording to the literature, laser shock peening (LSP) improves the surface quality of additively manufactured parts [26,27].

Effect of In-situ Laser Shock Peening and Tile Overlap on DensityAccording to the literature, laser shock peening (LSP) improves the surface quality of additively manufactured parts [26,27]. Therefore, an extra shot of the laser was applied after the solidification of each tile to improve the surface roughness of each layer and, thus, to have a more uniform layer interface and enhanced bonding. Figure 10 shows an improvement in the relative density of IN718 parts at different energy levels using the in-situ LSP strategy. Tile overlap is another variable that influences the density of the Area Printed parts. Figure 11 shows this effect is more pronounced at higher laser powers.)%2.01.5No peening Peening( y t i s o r o P1.00.50.00.160.18 0.22 0.20 Energy Density (kJ/cm2)0.24Figure 10.

Effect of in-situ laser shock peening on porosity of IN718 parts.)%( y t i s o r o P7 6 5 4 3 2 1 0Tile overlap: 200 μm 400 μm 600 μm5.09.0 Laser power (kW/cm2)13.0Figure 11. Effect of tile overlap on porosity of M300 maraging steel parts.3.2.Simulation resultsThe simulation was set up to understand the melting and solidification behavior of the LAPBF process. Since the simulation model was designed with very small meshes, laser pulse length transformation is applied based on analytical expression to approximate the diode energy densities. To compare similar melting behavior at different laser parameters, P × √𝑡 can be used1089as a constant for pulsed laser systems where P is laser power, and t is the pulse length [28].

To compare similar melting behavior at different laser parameters, P × √𝑡 can be used1089as a constant for pulsed laser systems where P is laser power, and t is the pulse length [28]. This enables lowering the pulse length significantly, thus lowering the simulation time.The simulation was carried out with a diode laser with approximate energy densities of ~0.09 kJ/cm2 and ~0.18 kJ/cm2 and YAG laser with ~30 J/cm2, 10 ns – 100 µs [pulse length Seurat proprietary]. The diode power was increased to 50 kW for the simulation, which lowered the pulse lengths to 0.5 ms and 1 ms, respectively. Figure 12 shows the state of the powder bed before the laser hits the powder bed. At low diode energy densities, some powder particles were partially melted. As the diode energy was increased to ~0.18 kJ/cm2, more powder particles were melted, and particle sintering and balling were dominant.Figure 12.

As the diode energy was increased to ~0.18 kJ/cm2, more powder particles were melted, and particle sintering and balling were dominant.Figure 12. Temperature distribution and powder sintering at the end of the diode laser heating with (a) 0.09 kJ/cm2 and (b) 0.18 kJ/ cm2.The YAG laser is applied to melt the powder layer completely. However, depending on the diode energy, the melted powder layer may or may not fuse to the substrate. Figure 13 presents the cross-sectional view of the melt pool at the end of the YAG laser for low energy diode. The applied lasers result in a high aspect ratio melt pool, and with a low energy diode, insufficient melting was observed, leading to high porosity. The melting and solidification with the YAG laser have also been discussed in previous work [29].Figure 13. Melt region at the end of the YAG laser applied after low energy diode in IN718.10904.

Melt region at the end of the YAG laser applied after low energy diode in IN718.10904. ConclusionsIn this study, large-area pulsed laser powder bed fusion parameters are optimized to achieve near-full-density parts (>99.9 % relative density) using Ni-based supper alloy 718 and M300 maraging steel. It is observed that the full-density area for IN718 is significantly narrower than that of M300. This is attributed to IN718 being prone to internal cracking at high energy densities because of the fast solidification and cyclic heating process. No inflection point was observed in the density plot of M300 samples by increasing the applied energy density. More experiments at higher energy densities and comparing the fully dense samples at various energies are yet to be carried out. Also, a computational model is developed to understand the multi-physics governing the process.

Also, a computational model is developed to understand the multi-physics governing the process. The simulation results showed that increasing the diode laser power raises the powder bed temperature and, as a result, increases the size of the melt region. The higher temperatures and larger melt regions ensure proper inter-layer bonding when applying the YAG laser. Also, it is observed that the shallow depth and high aspect ratio of the melt pool lead to a unidirectional solidification front extending along the build direction where grains grow epitaxially, and highly directional microstructures are created.References:[1]J. Li, J. Hu, Y. Zhu, X. Yu, M. Yu, H. Yang, Surface roughness control of root analogue dental implants fabricated using selective laser melting, Addit Manuf. 34 (2020) 101283. https://doi.org/10.1016/J.ADDMA.2020.101283.[2]S. Torres-Carrillo, H.R. Siller, C. Vila, C. López, C.A.

viewed PaperEXPERIMENTAL AND NUMERICAL ANALYSIS OF LATTICE STRUCTURES WITH DIFFERENT HETEROGENEITIESGuoying Dong, Yuchen Ding, Ladpha Teawdeswan, Chaoqian Luo, Kai YuDepartment of Mechanical Engineering, University of Colorado Denver, Denver, CO, 80204AbstractLattice structures with optimized material distributions can achieve unique mechanical properties such as high stiffness-to-weight ratio. However, the numerical analysis of the mechanical properties of heterogeneous lattice structures is challenging. In this research, three numerical approaches, including the beam element model, tetrahedral element model, and two- stage homogenization model, were used to predict the stiffness of lattice structures with different heterogeneities. Compression tests were conducted to evaluate the accuracy of the simulation results of each numerical approach. It was found that the accuracy of the numerical model varies with the increasing of heterogeneities.

It was found that the accuracy of the numerical model varies with the increasing of heterogeneities. The beam element model significantly underestimated the stiffness. The tetrahedral element model is the most accurate, but the computational cost is extremely higher than others. The results also indicated that, although the homogenization-based numerical model can substantially reduce the computational cost, the accuracy can be compromised due to the heterogeneity of lattice structures.Keywords: Additive Manufacturing, Finite Element Analysis, Heterogeneous Structure, Lattice StructureIntroductionLattice structures in additive manufacturing refer to structures with interconnected struts and nodes in a three-dimensional (3D) space [1]. The manufacturing, design, optimization, and simulation of lattice structures have been studied for over a decade. The lattice structure can be classified into homogeneous and heterogeneous structures based on homogeneity [2], as shown in Figure 1.

The lattice structure can be classified into homogeneous and heterogeneous structures based on homogeneity [2], as shown in Figure 1. If the relative density of the structure is uniform across the whole space, it is called homogeneous lattices. If the relative density has a gradient in the space, it is called heterogeneous lattices. Because the material distribution in the lattice structures can be optimized according to the functional requirements, heterogeneous lattice structures can exhibit better mechanical properties than homogeneous lattice structures. Most of the design and optimization methods in the literature for lattice structures are simulation-based. Finite element analysis (FEA) has been widely used to predict the mechanical properties of lattice structures.

Finite element analysis (FEA) has been widely used to predict the mechanical properties of lattice structures. The FEA methods can be categorized into three types: beam element model, solid element model, and homogenization model.2162Figure 1 An example of (a) homogeneous lattice and (b) heterogeneous lattice.Beam elements are prevalently used in FEA of lattice structures because of their low computational costs compared to solid elements [3]. Each lattice strut is meshed with several beam elements. The total degrees of freedom of the beam element model are much smaller than that of the solid element model. Timoshenko beam elements are widely used because the lattice struts are usually stout beams, in which shear deformation is important [4, 5]. However, there are some limitations in the beam element model. It is reported that the beam element cannot accurately model the joint of the lattice structure. It tends to underestimate the stiffness of the whole structure [6].

It is reported that the beam element cannot accurately model the joint of the lattice structure. It tends to underestimate the stiffness of the whole structure [6]. The beam element needs to be modified to consider the joint stiffening effect in the lattice structure [7]. Furthermore, it is known that the as-built lattice structure varies from the designed geometries [8]. The beam element model cannot reflect the actual geometry of the lattice structure. An alternative way is to vary the diameters of the beam elements according to the as-built strut thickness distribution [9]. Campoli et al. [10] used scanning electron microscopy and Gaussian distribution to determine the diameter of each beam.

Campoli et al. [10] used scanning electron microscopy and Gaussian distribution to determine the diameter of each beam. Then, the FE model was simulated several times to get the mean and standard deviation of the mechanical properties which means the lattice structure may have a range of mechanical properties due to the structural irregularity.In solid element models, tetrahedral [11] elements are mostly used to model the lattice structure. Because the element size should be smaller than the strut thickness to obtain a high- quality mesh, the number of elements in lattice structures can be extremely large. Therefore, the computational cost of the solid element model is significantly higher than the beam element model. Because solid elements can accurately model the geometry of the joint, the accuracy of the solid element model is higher than the beam element model, especially when the relative density of the lattice structure is high [12].

This is because the joints take more volume ratio, and the struts are stout when the relative density is high. Furthermore, solid elements can be directly built on the X- ray tomography model. The geometrical discrepancy caused by the unmelted powder, defects, and manufacturing errors can be accurately captured by this method. It can more accurately predict the mechanical properties influenced by the poor build quality of AM processes [13].The homogenization model can also be called a two-step FEA model. In the first step, the unit cell in the lattice structure is considered a representative volume element (RVE). The homogenization method is used to calculate the equivalent property of the RVE. In the second step, the lattice structure is considered a bulky part and meshed with hexahedral elements [14, 15]. Each hexahedral element represents one unit cell in the lattice structure. Therefore, the total number of elements in the homogenization model is very small.

Each hexahedral element represents one unit cell in the lattice structure. Therefore, the total number of elements in the homogenization model is very small. The advantage of the homogenization model is that it can dramatically reduce the computational cost if the equivalent property of the lattice cell can be found in a database [16]. However, the most obvious limitation of the homogenization2163model is that it can only work with periodic lattice structures. For randomized lattice structures, the homogenization method is not able to get the equivalent properties of the RVE.Compared to homogeneous lattice structures, heterogeneous ones bring more challenges to the simulation model. For the beam element model, the joints in the heterogeneous lattice structure are all different. It is difficult to study the mechanical behaviors of all the joints and incorporate them into the beam element model.

It is difficult to study the mechanical behaviors of all the joints and incorporate them into the beam element model. For the solid element model, the mesh size should be smaller than the minimum thickness of the strut, which will further increase the number of elements and the computational cost. For the homogenization model, the periodic boundary conditions are required for the RVE. It is originally proposed for homogeneous lattices only. Although the homogenization method has been used to model heterogeneous lattice structures [17], the accuracy of the simulation is still unknown. To figure out the accuracy and efficiency of different models in the simulation of heterogeneous lattice structures, in this research, the beam element model, solid element model, and homogenization model were used to simulate the lattice structure with both uniform and non-uniform density distributions. Experiments were conducted to evaluate the simulation models.

Experiments were conducted to evaluate the simulation models. The results were discussed to explain the discrepancy between the simulation results and the experimental results. Finally, conclusions were wrapped up and future research perspectives were pointed out.Methodology2.1 Design of the lattice structuresTo study the effect of the heterogeneity on the simulation models, the X-shape lattice structures were generated with the same overall relative density but different heterogeneities. Four types of structures were generated, and the parameters are shown in Table 1. The overall relative density of the lattice structure was set to 20%. The relative density of each unit cell was randomly generated. The upper and lower bounds of the unit cell were different in the lattice structures. L-1 was designed as a homogeneous lattice structure as the relative density of all the unit cells was 20%.

The upper and lower bounds of the unit cell were different in the lattice structures. L-1 was designed as a homogeneous lattice structure as the relative density of all the unit cells was 20%. L-4 was designed as the lattice structure with the highest heterogeneity.Table 1 The design parameters for the lattice structures studied in this research.NameL-1 L-2 L-3 L-4Overall relative density20%Maximum relative density 20% 25% 30% 35%Minimum relative density 20% 15% 10% 5%Unit cell size (mm)5Number of unit cells62164Figure 2 Relationship between the relative density and strut diameters.A group of lattice unit cells was generated with different radii. The relative density of each unit cell was measured. The relationship between the strut radius and the relative density was obtained by interpolation, as shown in Figure 2. Then, the randomly generated relative density of each unit cell was converted to the radius of the lattice strut.

Then, the randomly generated relative density of each unit cell was converted to the radius of the lattice strut. Finally, the geometrical models of four lattice structures were generated, as shown in Figure 3(a). The distribution of the radius of each unit cell is shown in Figure 3(b). L-1 has a uniform relative density, so the radius of unit cells is a constant. L-2 has a radius distribution from 0.48 mm to 0.64 mm. L-3 has a radius of distribution from 0.39 mm to 0.71 mm. L-4 has the maximum heterogeneity. The radius is from 0.28 mm to 0.78 mm. All the radii were randomly generated. The overall relative density of all the lattice structures is kept at 20%.Figure 3 (a) the geometrical model of the designed lattice structures, and (b) the radius distribution of the unit cells in the lattice structures.2.2 Simulation methodsThree FEA methods were used to simulate the design lattice structures: beam element model, solid element model, and homogenization model.

In the beam element model, the diameter of the beam was assigned individually. Two types of beam element models were studied in this2165research. One did not consider the joint influence of the lattice structure, as shown in Figure 4(a). The other increased the diameter of the beam element at the joint of the structure to consider the joint stiffening effect, as shown in Figure 4(b). In the solid element model, the lattice structure was meshed with coarse and fine tetrahedral elements, as shown in Figure 4(c). nTopology was used to generate the solid mesh model. The fine mesh use edge length 0.25mm and growth rate 2. The coarse mesh use edge length 1mm and growth rate 2. In the homogenization model, the relationship between the equivalent material properties and the relative densities of the X-shape lattice structure was obtained by the homogenization method [18], as shown in Figure 5.

Because the unit cell is symmetric, there are only three unknowns (C11, C12, C44) in the homogenized constitutive matrix. Then, the lattice structure was considered as a bulky part meshed by hexahedral elements. Each hexahedral element represents one unit cell. The material property assigned to the hexahedral element is the equivalent material property of the corresponding unit cell, as shown in Figure 4(d). A displacement load along Z-axis was applied on the top of the lattice structure, and the bottom of the lattice structure was fixed. The stiffness of the structure was predicted by the simulation model. It is defined by the ratio of the total reaction force along the Z-axis to the displacement of the top of the lattice. The simulation results would then be evaluated by experimental results.Figure 4 Example of the simulation models used in this research, (a) beam element model without considering the joint influence.

(b) beam element model with increased beam diameter at the joint areas, (c) solid element model with tetrahedral elements, and (d) homogenization model with hexahedral elements, of which the material properties are assigned by the equivalent material properties of the corresponding unit cell.Figure 5 The entries in the homogenized constitutive matrix vs. relative density of X-shape unit cell.21662.3 ExperimentsThe lattice structures were manufactured by a digital light processing (DLP) printer. The material is FAST abs-like resin provided by Siraya Tech. Each lattice structure was made by three times. A tensile specimen was designed to get the elastic modulus of the bulk material which was used in the simulation model. It was found that the elastic modulus of the material is 846 ± 49 MPa. Compression tests were conducted to get the force-displacement curve of the designed lattice structure, as shown in Figure 6.

It was found that the elastic modulus of the material is 846 ± 49 MPa. Compression tests were conducted to get the force-displacement curve of the designed lattice structure, as shown in Figure 6. MTS Insight 30 was used to do the compression test with a 30 kN load cell. The speed of the test was set to 1 mm/min.Figure 6 Compression test of the lattice structures.Results3.1 Experimental resultsThe results of the compression test are shown in Figure 7. It can be found from Figure 7(a) that the slope of the curve in the linear region varies among the lattice structures. The stiffness of the lattice structure is defined by the force increment divided by the displacement increment. The L-3 has the maximum stiffness in the linear region, but its force dropped earlier than the other structures. The overall force-displacement curve shown in Figure 6(b) indicates that the force of L-3 in the plateau region is the smallest among all the samples.

The overall force-displacement curve shown in Figure 6(b) indicates that the force of L-3 in the plateau region is the smallest among all the samples. To understand the relationship between the stiffness and the heterogeneity, the bar plot of the stiffness is shown in Figure 8. The height of the bar shows the average stiffness of the three samples. The error bar indicates the standard deviation of the samples. It was found that the homogeneous lattice structure L-1 has the minimum standard deviation of stiffness. The stiffness slightly dropped from 419 N/mm to 412 N/mm when the relative density changed from 20% (L-1 sample) to a random distribution between 15% and 25% (L-2 sample). However, the stiffness of the structure dramatically increased to 547 N/mm when the relative density distribution was between 10% and 30% (L-3 sample). Then, the stiffness drops again to 532 N/mm when the relative density distribution is between 5% and 35% (L-4 sample).

Then, the stiffness drops again to 532 N/mm when the relative density distribution is between 5% and 35% (L-4 sample). From samples L-1 to L-4, the heterogeneity of the lattice structure was increased. But the stiffness didn’t increase monotonously with the heterogeneity. The experimental results will be used to evaluate the accuracy of different simulation models studied in this research.2167Figure 7 The force-displacement curve of the lattice structures under compression test.Figure 8 The stiffness of the lattice structures with different heterogeneities.3.2 Simulation resultsThe computational costs of the simulation models are related to the number of elements in the model. Table 2 shows the number of elements in each model. It can be found that the homogenization model has the minimum number of elements, which is equal to the number of unit cells in the lattice structure. The beam element model meshed each lattice strut with one element.

The beam element model meshed each lattice strut with one element. Thus, the total number of struts in the lattice structure determined the number of beam elements. The joint stiffening effect was considered in a modified beam element model. In this model, each lattice structure was meshed by 10 beam elements. The radius of the beam elements at both ends of the strut was increased 10 times to compensate for the underestimation of the joint stiffness. The total number of the elements was equal to 10 times the number of the strut. Furthermore, two tetrahedral element models were used in this research. The mesh sizes of these two models are different. In the coarse model, the number of elements is around 200,000.

It was found that the results in different simulation models vary a lot. The beam element model without the consideration of the joint gave the minimum stiffness prediction. The coarse tetrahedral element model obtained the maximum stiffness prediction. To evaluate the accuracy of the simulation models, the simulation results were compared to experimental results, as shown in Figure 9. The actual stiffness of the structure is around 3 times the prediction by the beam element model without the consideration of the joint. Even if the size of the joint in the beam element model is increased 10 times, the predicted stiffness is only half of the experimental result. Therefore, the accuracy of the beam element model is the lowest among all the simulation models.The tetrahedral element models can more accurately predict the stiffness of the structure. However, the coarse mesh model would overestimate the stiffness of the structure by around 30%.

However, the coarse mesh model would overestimate the stiffness of the structure by around 30%. Although the computational cost of the fine tetrahedral element model is significantly higher than other simulation models. It can most accurately predict the stiffness of lattice structures among all the simulation models. The homogenization model also underestimated the stiffness of lattice structures. The difference between the homogenization result and the experimental result is around 10%, which is the second most accurate prediction among all the simulation models. It should be noted that the homogenization model has the lowest computational cost.2169Figure 9 The comparison of stiffness predicted by different simulation models to the experimental result.DiscussionIt is found from the experimental results that the lattice structures with the same weight but different heterogeneities exhibit different mechanical properties.

The stiffness changes in the beam element model with joint stiffened, fine tetrahedral element model, homogenization model, and the experimental results are shown in Figure 10. The green line represents the change in the experimental result. From L-1 to L-4, the heterogeneity of the lattice structure was increasing. The stiffness slightly dropped from 419 N/mm to 412 N/mm and then increased to 547 N/mm dramatically. Because the relative density distribution is randomly generated, the change of the mechanical property is also random. The experimental result is a benchmark to evaluate the simulation models.Figure 10 The changes in the stiffness of lattice structures with different heterogeneities in the simulation models and the experimental results2170The result of the beam element model didn’t change with the increase in heterogeneity. The maximum change of the stiffness in the beam element model is less than 2%.

The maximum change of the stiffness in the beam element model is less than 2%. It was also found that the stiffness of L-3 predicted by the beam element model had the minimum value in the four structures. However, the experimental result showed that the stiffness of L-3 is the highest in all the structures, which contradicts the beam element model. It can be concluded that the beam element model cannot capture the change of stiffness with the change of heterogeneity. It also predicted the most inaccurate result. It should be noted that the lattice structure studied in this research is a bending dominant type. The joint influence on the mechanical properties of the bending dominant lattice structure is significant. The beam element cannot accurately model bending dominant lattice structures without precisely considering the joint influence. However, the joints in the heterogeneous lattice structure are not consistent.

However, the joints in the heterogeneous lattice structure are not consistent. The joint geometry is determined by the strut thickness connected to the joint. A model that can capture the joint influence in the heterogeneous lattice structure is needed to improve the accuracy of the beam element model.The stiffness predicted by the tetrahedral element model changes the most among the three simulation models shown in Figure 10. It is the only simulation model that predicted a lower stiffness of L-2 than that of L-1. Then, the stiffness predicted by the tetrahedral element model increased when the heterogeneity increased from L-2 to L-3. It means that the tetrahedral element model can indicate the stiffness change with the heterogeneity change. However, the predicted stiffness kept increasing when the heterogeneity increased to L-4, which is different from the experimental result.

However, the predicted stiffness kept increasing when the heterogeneity increased to L-4, which is different from the experimental result. Although there is little discrepancy between the experimental results and the tetrahedral element prediction, it is the most accurate one among all the simulation models. One of the reasons is that the tetrahedral element considers the lattice structure as a general solid part. There are no other assumptions like the beam element model and the homogenization model. But the complex geometry of the lattice structure leads to a high computational cost of the tetrahedral element model. It also has a high requirement for mesh quality. A coarse mesh model may overestimate the stiffness of the lattice structure. Mesh sensitivity analysis is needed to avoid the inaccuracy caused by the mesh quality.Because the homogenized material property of the lattice unit cell is obtained from a database.

Mesh sensitivity analysis is needed to avoid the inaccuracy caused by the mesh quality.Because the homogenized material property of the lattice unit cell is obtained from a database. The homogenization model has the lowest computation cost among all the simulation models. It was found in Figure 10 that the stiffness prediction of the homogenization model increased with the increase of the heterogeneity. It didn’t reflect the actual stiffness change with the heterogeneity in the experimental result. The possible reason is that, in the homogenization method, the material property is obtained based on a periodic boundary condition on the lattice unit cell. It was developed for homogeneous lattice structures. However, this assumption is not valid for heterogeneous lattice structures. The two-step homogenization FEA model gets inaccurate with the increase of heterogeneity.

However, this assumption is not valid for heterogeneous lattice structures. The two-step homogenization FEA model gets inaccurate with the increase of heterogeneity. Nevertheless, the accuracy of the homogenization model is much higher than the beam element model, and the computational cost is significantly lower than the tetrahedral element model. Therefore, this method has been used to model heterogeneous materials and structures in the literature. Future research is needed to modify the homogenization model and increase the accuracy for highly heterogeneous structures.2171Heterogeneous lattice structures can be used in many applications that require a high stiffness-to-weight ratio, energy absorption rate, or specific stiffness distribution. Therefore, an accurate simulation model for heterogeneous lattice structures is important to the design process.

Therefore, an accurate simulation model for heterogeneous lattice structures is important to the design process. It is found from this research that there are still gaps between existing strategies and the actual mechanical properties of heterogeneous lattice structures. The computational cost of the fine tetrahedral element model is too high. The homogenization model needs improvements to deal with heterogeneous problems. Furthermore, nonlinear analysis of heterogeneous lattice structures is more challenging for simulation models. It will dramatically increase the computational cost of the tetrahedral element model. Currently, the homogenization model cannot deal with nonlinear problems. The nonlinear mechanical property of the lattice unit cell should be obtained in the first step. The feasibility of the two-step FEA model needs to be validated in future research.ConclusionIn this research, the stiffness of the lattice structures with different heterogeneities was studied.

The feasibility of the two-step FEA model needs to be validated in future research.ConclusionIn this research, the stiffness of the lattice structures with different heterogeneities was studied. 3 types of simulation models were used to predict the stiffness. Compression tests were conducted to evaluate the accuracy of each simulation model. It was found from the experimental result that the stiffness does change monotonously with the heterogeneities. It was also found that the results from different simulation models vary a lot from each other. The beam element model has the lowest stiffness prediction, mainly because the joint stiffness is underestimated. Even if the radius of the beam elements at the joint increased 10 times, the prediction was still half of the experimental result. The tetrahedral element model has the highest accuracy, but the computational cost is significantly higher than other simulation models. Furthermore, it is sensitive to the mesh quality.

The tetrahedral element model has the highest accuracy, but the computational cost is significantly higher than other simulation models. Furthermore, it is sensitive to the mesh quality. If the mesh is coarse, the tetrahedral element model will overestimate the stiffness by 30%. The homogenization model has a balance between accuracy and computational cost. The stiffness prediction is about 10%-20% lower than the experimental value.The influence of heterogeneity on the accuracy of the simulation model was also studied. It was found that the beam element model cannot reflect the stiffness change caused by the heterogeneity of the lattice structure. The homogenization model results showed that the stiffness increased with the increase of heterogeneity. However, it was inconsistent with the experimental result. The tetrahedral element model can most accurately predict the stiffness change caused by the increased heterogeneity.

However, it was inconsistent with the experimental result. The tetrahedral element model can most accurately predict the stiffness change caused by the increased heterogeneity. But it also lost the accuracy when the heterogeneity kept increasing. Overall, all the simulation models cannot accurately predict the stiffness of lattice structures with high heterogeneity. Future research can focus on the improvement of the homogenization model because of the low computational cost and relatively high accuracy. The nonlinear analysis should also be studied for the homogenization method.AcknowledgmentThis research is supported by funding from the University of Colorado Denver, College ofEngineering, Design, and Computing.21721.2.3.4.5.6.7.8.9.10.11.12.13.14.15.16.17.18.ReferenceDong, G., Y. Tang, and Y.F. Zhao, A survey of modeling of lattice structures fabricated by additive manufacturing. Journal of Mechanical Design, 2017. 139(10): p. 100906.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperExperimental and Numerical Investigations on Dynamic Mechanical Properties of TPMS StructuresDeepak Kumar Pokkalla1*, Brandon Turner White2, Jier Wang3, Ryan Spencer2, Ajit Panesar3, Seokpum Kim1 and Uday Vaidya21Manufacturing Science Division, Oak Ridge National Laboratory, TN 37932, USA 2Tickle College of Engineering, University of Tennessee, Knoxville, TN, 37916, USA 3Department of Aeronautics, Imperial College London, SW7 2AZ, UKCorresponding author: pokkallad@ornl.govAbstract Triply Periodic Minimal Surface (TPMS) lattice structures have been of increasing interest due to their light weighting, enhanced mechanical properties, and energy absorption characteristics for automotive and biomedical applications.

With the advent of additive manufacturing and geometric modeling software, TPMS lattices with complex geometries can be realized. In this work, TPMS lattice structures were fabricated with PLA using fused filament fabrication (FFF) and their dynamic properties are characterized through drop tower experiments. Although lightweight TPMS lattices are beneficial for their impact absorption capability, most of the existing works are limited to quasi-static compression, and dynamic impact tests are rarely performed. The current study investigates the stress-strain and energy absorption characteristics of TPMS lattices through drop tower testing and numerical modeling. Finite element modeling for TPMS lattices is carried out to validate the experimental responses.

Finite element modeling for TPMS lattices is carried out to validate the experimental responses. The mechanical properties, deformation, and failure mechanisms of TPMS lattices under dynamic impact are summarized for potential future applications.IntroductionIn recent years, lattice structures have gained popularity due to their light weighting, enhanced mechanical properties such as energy absorption, and thermal management capabilities for automotive, aerospace, and biomedical applications [1-4]. With the advent of additive manufacturing, a range of lattice structures in custom and complex shapes can be realized. Incorporating these lattice structures into traditional components to reduce weight is becoming increasingly feasible with readily available software like nTopology, Altair, and Materialise. For instance, lattice structures can be used in orthopedic implants as scaffolds for bone and tissue regeneration [5-6].

For instance, lattice structures can be used in orthopedic implants as scaffolds for bone and tissue regeneration [5-6]. Lattice cores can be used in industrial heat exchangers to improve their efficiency [7].The selection of a suitable lattice for a given application is often a cumbersome task given the vast array of lattice structures, despite classifying them majorly into volumetric or surface-based and periodic, non-periodic, or stochastic lattices [8]. The lattice type, size, material, and manufacturing process can drastically change the mechanical or thermal response of the overall lattice-based structural component [3-4,9]. Hence, it is essential to choose the appropriate lattice structure with geometrical features and material to achieve desired mechanical properties for industrial applications.This manuscript has been authored in part by UT-Battelle, LLC, under contract DE-AC05-00OR22725 with the US Department of Energy (DOE).

The US government retains and the publisher, by accepting the article for publication, acknowledges that the US government retains a nonexclusive, paid-up, irrevocable, worldwide license to publish or reproduce the published form of this manuscript, or allow others to do so, for US government purposes. DOE will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan). 1876Among various lattice structures, the triply periodic minimum surface (TPMS) lattices have been gaining interest due to their high strength and energy absorption characteristics [10]. TPMS lattices are also known for carrying more loads and producing smoother crushing because of more uniform stress distributions [11].

TPMS lattices are also known for carrying more loads and producing smoother crushing because of more uniform stress distributions [11]. Despite the advantages of the TPMS lattice structures, such as high energy absorption, most of the existing studies for composite TPMS lattices are limited to studying the elastic or quasi-static mechanical responses [3]. Various functional gradients such as density and cell size have also been introduced to evaluate the effect of tailored designs on the compression responses of TPMS lattices [4]. To this end, this work focuses on evaluating the dynamic mechanical responses of the polymer TPMS lattices, specifically gyroid and diamond, for different impact velocities.Finite element analysis is commonly used for the numerical evaluation of the mechanical properties of lattices structures [12-14].

Several finite element-based optimization frameworks have also been developed for the design of lattice structures to achieve prescribed mechanical responses over large strains [15-16]. In addition, several different strategies can be used to model these TPMS lattices such as shell, solid, and voxel models. The choice of an appropriate model depends on the accuracy, computational efficiency, ease of modeling, and is often considered as a trade-off from one another. For TPMS lattices, voxel models with hexahedral finite element meshes have shown potential to provide accurate mechanical property predictions with high computational efficiency [3]. Hence, gyroid and diamond lattice structures were modelled through voxels for the numerical simulations in this work.Fabrication and ExperimentsThe TPMS lattices are surface-based lattice topologies that can be described using mathematical equations.

Gyroid (G) and diamond (D) lattices can be generated with the prescribed number of unit cells and volume fraction from the following equations.𝐶𝐺 = cos(𝑋) ∗ sin(𝑌) + cos(𝑌) ∗ sin(𝑍) + cos(𝑍) ∗ sin(𝑋) − 𝑡𝐶𝐷 = sin(𝑋) sin(𝑌) sin(𝑍) + sin(𝑋) cos(𝑌) cos(𝑍) + cos(𝑋) sin(𝑌) cos(𝑍) + cos(𝑋) cos(𝑌) sin(𝑍) − 𝑡where 𝑋, 𝑌, and 𝑍 are given by 2∗𝜋∗𝑘 with 𝑘 and 𝐿 prescribing the number of unit cells and length of the structure in each direction. The volume fraction of the lattice structure can be tuned using the parameter 𝑡. The change in volume fraction within the unit cells could result in vastly different mechanical responses of these lattice structures.𝐿In this work, gyroid and diamond lattices were chosen and their mechanical responses under dynamic loading were investigated.

The geometries for these lattices with 50% volume fraction in a configuration of 7 x 7 x 7 that fit into a 40 mm side cube were generated using the above mathematical equations and the corresponding STL files were generated as shown in Fig. 1. The STL files were sliced in Ultimaker Cura with an infill density of 100%, a standard quality layer height of 0.2 mm, a printing temperature of 210 ℃, and a print speed of 80.0 mm/s. Four specimens of gyroid and three specimens of diamond lattice were manufactured using PLA with an Ender 3 Pro desktop 3D printer. The printed gyroid and diamond lattice structures are depicted in Fig.

The printed gyroid and diamond lattice structures are depicted in Fig. 2.1877Figure 1: As designed TPMS lattice structures (gyroid and diamond) with a configuration of 7 x 7 x 7- unit cells in a side length of 40 mm and 50% volume fraction.Figure 2: As printed TPMS lattice structures using PLA with an Ender 3 Pro desktop 3D printer.The total mass of the printed lattice blocks was measured and averaged to obtain a total mass of 35.74 g and 34.79 g for gyroid and diamond lattices, respectively. The side length of the lattice cubes was also measured to be approximately 40 mm. These fabricated specimens were subjected to dynamic impact testing using a drop tower shown in Fig. 3.1878Figure 3: (a) Intron CEAST 9340 Drop Tower, (b) Setup with 22kN tup, six 5 kg load cells, and60 mm flat tup head, and (c) Setup with tup height set.The Instron CEAST 9340 drop tower (refer Fig. 3(a)) was used for drop tower experiments in this work.

3(a)) was used for drop tower experiments in this work. The CEAST data acquisition system measures the force (N), energy absorption (J), and displacement (mm) using the kinetic energy equation based on the predetermined drop height and velocity. A 22kn tup head was used for the data acquisition with a 60 mm Ø flat tup head, and the overall mass of impact was set to 33.215 kg, which has a base standard weight tup holder (7510.021) mass of 3.215kg with an additional mass of 30kg as shown in Fig. 3(b). The data acquisition system used was the CEAST DAS 64k - SC High Speed Data Acquisition Unit (7191.000). The parameters for the experiment were set to a frequency of 2000kHz and 64,000 data points with a working force range of 31kN.

The parameters for the experiment were set to a frequency of 2000kHz and 64,000 data points with a working force range of 31kN. Each experiment complied with ASTM D7136 standard test type.For the experimental setup, the adjustable height stand (7520.035) below the 3D structure was adjusted so that the lowest point of the flat tup-head could only reach the third from the bottom layer of each lattice block so that no damage could occur to that region and avoids hitting the metallic fixture. Before the test could initiate the tup-head’s position was adjusted to slightly above the top of the lattice structure, as seen in Fig. 2(c), and the laser for when the data acquisition begins is set to that height. The 3D structure must be centered relative to the tup-head, and once centered, the drop test can be initiated. Each specimen test was conducted at a constant velocity of 4.5 m/s and impact height of 1032 mm.

Each specimen test was conducted at a constant velocity of 4.5 m/s and impact height of 1032 mm. The 30kg constant mass was determined to induce sufficient deformation in the specimens based on the preliminary tests conducted with weights of 5kg, 10kg, 20kg, 25kg, and 30kg. The samples in each preliminary test conducted were evaluated for any visual damage along with their force-time graphs. Upon these preliminary experiments, a constant weight of 30 kg is used subsequently for testing the gyroid and diamond lattice structures.Numerical SimulationsModelling and finite element simulations are beneficial for rapidly evaluating the mechanical properties of various TPMS lattice structures with different volume fractions and geometrical features. In this work, finite element analysis on lattice structures was performed using ABAQUS/Explicit solver.

In this work, finite element analysis on lattice structures was performed using ABAQUS/Explicit solver. There are several different meshing strategies that can be explored for surface-based lattice structures such as shell, solid, and voxel models. Depending upon the geometry and the volume fractions, these models provide a trade-off between the accuracy, modelling ease and efficiency, and data management.1879In this work, a gyroid lattice structure with 7x7x7 unit cells (side length of 40 mm) and 50% volume fraction is meshed using a voxel model following the procedure presented in [3] and as shown in Fig. 4. The gyroid lattice was placed on a lower fixed rigid plate and the impactor is modeled as a top rigid plate. The top impactor with a weight of 30 kg hits the lattice structure at a velocity of 4.5 m/s. A nonlinear elastic-plastic material model calibrated from the uniaxial test data on PLA [17] is used for material model here.

A nonlinear elastic-plastic material model calibrated from the uniaxial test data on PLA [17] is used for material model here. The elastic modulus of the PLA is determined to be 2554.7 MPa and an elastic Poisson’s ratio of 0.33. A non-linear isotropic hardening is used in the plastic region. The rate- dependency of the PLA is ignored in this current study as the variation in elastic modulus for small impact velocities up to 4.5 m/s is found be small from the dynamic mechanical analysis.

The rate- dependency of the PLA is ignored in this current study as the variation in elastic modulus for small impact velocities up to 4.5 m/s is found be small from the dynamic mechanical analysis. Modelling the strain-rate dependency effects for accurate estimation of the mechanical properties of the lattice structures subjected to high-rate impacts is part of our ongoing work.Figure 4: (a) Gyroid lattice structure with 7x7x7 unit cells and 50% volume fraction, (b) Hexahedral finite element mesh with C3D8R elements, (c-d) Boundary conditions – fixed bottom rigid plate and top plate impacting at a velocity of 4.5 m/s.The modeling approach for the voxel model is adopted from [3], where surface-based geometry of the TPMS lattice is converted into a 3D array that was mapped directly to a hexahedral mesh with C3D8R elements. A 3D uniform mesh is embedded to encompass the lattice geometry where the voxels mapped to the solid material and voids can be easily distinguished.

A 3D uniform mesh is embedded to encompass the lattice geometry where the voxels mapped to the solid material and voids can be easily distinguished. The errors associated with the voxel-based mesh discretization were alleviated by performing a mesh refinement study and choosing the appropriate resolution that provides converged mechanical responses. Python scripts were developed to generate the voxel models and assign the material properties, boundary conditions, and analysis steps automatically once the type of lattice, volume fraction, and voxel resolution are chosen. Finite element analyses on gyroid lattice structures with various numbers of unit cells and volume fractions were performed.

Finite element analyses on gyroid lattice structures with various numbers of unit cells and volume fractions were performed. The numerical model developed will be compared with the experiments in the next section for validation, and the next steps are outlined.1880Results and DiscussionIn this section, the experimental and numerical results on the dynamic mechanical properties of the gyroid and diamond TPMS lattice structures are presented. The drop tower tests were first performed on a set of gyroid samples with different drop weights to determine the weight that would induce large strains. The samples selected consist of 7x7x7 unit cells and the overall dimensions of the cube are 40 mm x 40 mm x 40 mm. During the drop tower experiments, the force is measured from the sensors at the top of the tup. The corresponding stress is determined by dividing the measured force by the top cross-sectional area of 40 mm x 40 mm.

The corresponding stress is determined by dividing the measured force by the top cross-sectional area of 40 mm x 40 mm. Similarly, the effective strain is measured as the ratio of the top displacement to the original height of the lattice as 40 mm.Preliminary investigations on gyroid samples with various weights indicated that a drop weight of 30 kg would induce large strains in the specimens. Hence, the drop tower impact tests are performed on four gyroid specimens with the same drop weight of 30 kg and the corresponding stress-strain curves were extracted as shown in Fig. 5(a). These lattices boast enhanced energy absorption capabilities and thus, the energy absorbed with time of the drop tower impact tests were plotted in Fig. 5(b).

5(a). These lattices boast enhanced energy absorption capabilities and thus, the energy absorbed with time of the drop tower impact tests were plotted in Fig. 5(b). All four tests showed good repeatability and the gyroid lattice is shown to absorb up to 250J of energy.Figure 5: Dynamic mechanical response of gyroid lattices for an impact velocity of 4.5 m/s and drop weight of 30 kg: (a) Stress-strain curves, and (b) Energy absorption characteristics of four specimens with 50% volume fraction.Next, a diamond lattice structure was chosen, and the drop tower tests were repeated on it with the same 30kg drop weight. The stress-strain curves and the energy absorption plots are depicted in Fig. 6(a) and (b), respectively. It is evident that the 1st peak stress of 16.5 MPa occurs around 10% strain for diamond lattice whereas for gyroid the 1st peak stress value is 10.5 MPa. Comparing the energy absorption plots in Fig. 5(b) and Fig.

Comparing the energy absorption plots in Fig. 5(b) and Fig. 6(b) reveals that the energy absorption observed under dynamic impact in case of diamond lattices are more due to the higher peaks and valleys in the stress-strain curves.1881Figure 6: Dynamic mechanical response of diamond lattices for an impact velocity of 4.5 m/s and drop weight of 30 kg: (a) Stress-strain curves, and (b) Energy absorption characteristics of four specimens with 50% volume fraction.Dynamic explicit finite element simulations were performed on voxel model for the gyroid lattice structure and compared with the average stress-strain response obtained from the experiments as shown in Fig. 7. The dynamic explicit numerical simulation with mass scaling predicted the dynamic elastic modulus accurately, however, overpredicted the transition to the nonlinear regime.

7. The dynamic explicit numerical simulation with mass scaling predicted the dynamic elastic modulus accurately, however, overpredicted the transition to the nonlinear regime. The vowelized model is overpredicting the stiffness as the voxels corresponding to void region are modelled with low elastic modulus instead of removing them completely. The von Mises stress contours for the gyroid unit cells during the dynamic impact are shown in Fig. 8. The compression and high stresses occur at the critical locations within the unit cells and propagate throughout the geometry with further impact.

8. The compression and high stresses occur at the critical locations within the unit cells and propagate throughout the geometry with further impact. This indicates that more elements are subjected to higher stress with impact, until final failure occurs due to damage.Figure 7: Comparison of numerical stress-strain response with average response from the drop tower experiments for an impact velocity of 4.5 m/s and drop weight of 30 kg.1882Figure 8: Resultant von Mises stress contours for the gyroid lattice structure subjected to impact at 4.5 m/s velocity with a 30 kg impactor.ConclusionIn this paper, the mechanical properties of two TPMS lattice structures, namely gyroid and diamond under dynamic compression, were systematically investigated. The lattice structures were first printed using PLA in a desktop printer with enough unit cells within overall dimensions prescribed by the drop tower experimental setup.

The lattice structures were first printed using PLA in a desktop printer with enough unit cells within overall dimensions prescribed by the drop tower experimental setup. The response of the lattice structures under different drop weights was first studied and 30 kg drop weight was selected for further investigation. The stress-strain and energy absorption characteristics of gyroid and diamond lattices under drop tower impact with the same drop weight were experimentally determined. It was evident that diamond lattices undergo higher 1st peak stress than gyroid lattices under dynamic impact and thus have better energy absorption capability overall. A voxelized dynamic explicit numerical model was further developed to evaluate the mechanical response of the gyroid lattice. The comparison of numerical results with experiments revealed that the numerical predictions lead to stiffer response.

The comparison of numerical results with experiments revealed that the numerical predictions lead to stiffer response. One possible reason is that the voxels corresponding to void are not removed in the numerical model, instead small modulus values were assumed. Further development of the numerical model to match the experimental results and in-depth investigation of damage and failure mechanisms in TPMS lattice structures is part of our ongoing work.AcknowledgmentsThe research is in part supported by the US Department of Energy (DOE), Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle LLC.1883References1. Ashby, M.F., Evans, T., Fleck, N.A., Hutchinson, J.W., Wadley, H.N.G. and Gibson, L.J.,2000. Metal foams: a design guide. Elsevier.2. Al-Ketan, O., Lee, D.W., Rowshan, R. and Al-Rub, R.K.A., 2020.

PERIMENTAL AND NUMERICAL STUDY ON THE FLOW OF FINE POWDERS FROM SMALL-SCALE HOPPERS APPLIED TO SLS MULTI-MATERIAL DEPOSITION – PART IJames Santosa, Dejun Jing and Suman Das Department of Mechanical Engineering The University of Michigan Ann Arbor, MI 48109-2125AbstractWe present experimental guidelines for the delivery of powders under 100(cid:181)m throughhopper-nozzle orifice diameters on the order of 1mm. Small-scale hoppers will be incorporated into an SLS powder deposition system for creating thin layers of multiple powdered materials in a patterned bed. This is a preliminary investigation on the flow behavior for selected orifice diameters and particle sizes under gravity or low pressure-assisted flow conditions.

This is a preliminary investigation on the flow behavior for selected orifice diameters and particle sizes under gravity or low pressure-assisted flow conditions. A method for numerically modeling the gas-particle behavior in hopper-nozzles is presented and conditions for achieving continuous mass flow rates are demonstrated.INTRODUCTIONThe development of SFF techniques for producing a new class of artifacts with spatially-varying structure and multi-functional characteristics is in part dependent upon the ability to deposit and consolidate multiple materials. In particular, the selective laser sintering (SLS) process is well-suited to the incorporation of multiple powdered materials. The SLS process presently uses a roller device to sweep thin layers of a single powdered material across the build area.

The SLS process presently uses a roller device to sweep thin layers of a single powdered material across the build area. It has been proposed that this roller device be replaced by an array of hopper-nozzles that discretely deposit lines, dots and regions of multiple powdered material in a patterned bed.The designated name (cid:147)hopper-nozzle(cid:148) refers to the design of experimental nozzles based on existing hopper theory. In the chemical and process industries, hoppers have been inexpensively designed to store, discharge bulk solids, and eliminate undesirable flow instabilities (i.e. arching, rat-holing, and oscillatory flow). Unfortunately, difficult hopper and powder sizes are avoided due to the lack of fundamental understanding of flow phenomena.

arching, rat-holing, and oscillatory flow). Unfortunately, difficult hopper and powder sizes are avoided due to the lack of fundamental understanding of flow phenomena. This can be attributed to the complex interaction of granular solid and interstitial fluid that plays a large role in delivery through small orifice diameters.BACKGROUNDGranular bulk solids, like powder, exhibit characteristics unique from any other state. Thegranular mass is generally an amorphous, random-packing of particles influenced by the interstitial fluid occupying its voids. Unlike a fluid, the pressure under a vertical column of granular material is independent of its height, which makes a constant flow rate of material possible. Most flow theory has targeted particles above 500(cid:181)m that are cohesionless (fair to free- flowing in nature) and insensitive to drag. Generally, particles below 50(cid:181)m are small enough to be subject to cohesive effects.

Generally, particles below 50(cid:181)m are small enough to be subject to cohesive effects. Those used with SLS are typically 0.1 to 100(cid:181)m in diameter, and are more susceptible to air resistance when passing through the hopper orifice.620Evidently, the geometry of a hopper design is essential in achieving desired flow properties of the discharged material. The delivery of particles may happen under core flow (first in-last out) or the more desirable mass flow (first in-first out). For the design of a conical hopper, the hopper half-angle, α, and the orifice diameter, D0 are the most important dimensions to be considered in preventing arching (bridging of particles that blocks flow) and rat-holing (an empty tunnel over the orifice due to core-flow).

Jing and Li [9] provide useful information on hoppers for fine powders, stating an effective hopper half-angle of 10.2(cid:176) or less for maintaining mass flow.One of the most frequently used theories on flow from hoppers is W. A. Beverloo(cid:146)s modified correlation [2] for mass flow rate in conical hoppers. While there are other similar theories for mass flow rate, e.g. the hourglass theory, these do not take into account the effects of particle diameter.

While there are other similar theories for mass flow rate, e.g. the hourglass theory, these do not take into account the effects of particle diameter. The modified Beverloo correlation is= CGsρ bg+  dp dz  01 ρ b(D0−kdp)2/5(2)where Gs= mass flow rate of particles (g/s)C =empirical constant (usually .58) dependent on hopper angle and angle of internal friction bulk density of the powder (g/m3) gravitational constant of acceleration (9.8m/s2) hopper orifice diameter (m) empirical constant for particle shape (1.6 for spherical particles)ρB= g = D0= k = dp = mean particle diameter (m).The term 1/ρB (dp/dz)0 is a modification of the gravitational acceleration term to include the interstitial pressure gradient. Beverloo(cid:146)s correlation states that as particle size decreases, there is an increase in the mass flow rate.

Beverloo(cid:146)s correlation states that as particle size decreases, there is an increase in the mass flow rate. Empirical results agree for particle sizes down to 500(cid:181)m, where actual mass flow rate takes on much lower values than predicted.The intent of this study is to determine whether a correlation for SLS powder sizes may be extended from Beverloo(cid:146)s correlation with low-pressure assistance. Research on granular flow is still in the exploratory phase, and the time-dependent behavior of a bulk powder material may not yet be based on characteristics of the constituent particles [6]. Progress in this area must depend upon uncovering new correlations between the observed behavior and measured particle characteristics [16].

Progress in this area must depend upon uncovering new correlations between the observed behavior and measured particle characteristics [16]. With this in mind, the study aims to provide experimental results on the use of hoppers with fine powders, the design of the test apparatus and numerical model, and results of preliminary mass flow rate experiments conducted on a number of particle sizes and hopper- nozzles.621Figure 1. Pipette tips, pipets and drawn pipets are used as test hopper-nozzles.Figure 2. Pipette Figure 3. A high resolution image of the cut tip dimensions. orifice at approximately 40 X magnification.EXPERIMENTSHopper-Nozzle DesignTypically used for repetitive liquid dispensing, pipette tips were an inexpensive andpractical choice for producing multiple, adjustable nozzles. The 1-200(cid:181)L Uni-Tip' from BioPlas Inc., Figure 1 (a), is a siliconized polypropylene pipette tip with polished internal surfaces to help eliminate sample residue.

The 1-200(cid:181)L Uni-Tip' from BioPlas Inc., Figure 1 (a), is a siliconized polypropylene pipette tip with polished internal surfaces to help eliminate sample residue. The orifices of pipette tips are typically in the range of 0.75mm with a half-angle of 5.95(cid:176). They are designed for precise volumetric delivery; however there is no guarantee on the uniformity of the orifice diameters. A pack of 48 tips may have deviation up to 0.1mm.Because the tips are polypropylene, they were fairly easy to trim with a utility knife undera magnifying glass. Images taken with a high-resolution digital camera (Figure 3) were used to make diameter adjustments and verify the quality of the cut surface. Tips with visible striations, residual material around the inner diameter, or irregular cross-sections were re-cut or eliminated. Diameters were calibrated with images taken from a 5cm long, 5mm nominal I.D. glass tube next to a measurement scale.

Diameters were calibrated with images taken from a 5cm long, 5mm nominal I.D. glass tube next to a measurement scale. A digital caliper is then used to measure ten different diameters along the glass tube from a printout of the cross-section. A measured diameter of 5.001+/-0.027mm indicated that using the high resolution digital camera is an accurate scheme for measurement. Images from the cut pipette tips were measured in the same manner, with magnification between 30X and 60X. The nozzles were re-cut until they were within 0.005mm of the nominal diameter. Fifteen orifice diameters from 0.75mm to 2.5mm in steps of 0.125mm were created for testing. As an alternative to the polypropylene tips, Pyrex 5mL glass pipets, Figure 1(b), were diamond- ground to diameters 1.0-2.0mm in steps of 0.1mm. The graduated markings on the pipets are used to observe the pressure changes with respect to the height of the powder column.

The graduated markings on the pipets are used to observe the pressure changes with respect to the height of the powder column. Also by heating and drawing pipets, smaller diameters in the 0.5 to 1.0mm range were produced.Mass Flow Rate ExperimentsThe test particles were high-quality soda-lime glass beads, verified by a distribution histogram as 90% within the specified U.S. sieve mesh sizes and 90% spherical. Although622uniform particles of an identical diameter are considered the most desirable, it is believed that SLS powders are better represented by small ranges of particle size within 10-100 micron. The particles have diameter ranges d=10-25, 38-45, 45-53, 63-75, and 90-106 (cid:181)m.

The particles have diameter ranges d=10-25, 38-45, 45-53, 63-75, and 90-106 (cid:181)m. A majority of techniques for bulk materials testing encourage the use of a batch at least a few kilograms [16] however a few grams of each powder are acquired considering the particles are a sufficiently representative sample.The A&D GF-200 Precision Balance, a 0-200g, 0.001g resolution digital balance with RS-232C serial interface was used for sampling. The balance has Windows Communications Tools' for easy data transfer to Windows' 95/98. In its most rapid response mode, the balance has a 10Hz sampling frequency. Data can be imported into Microsoft Excel' or stored as unformatted data for other applications. In addition to mass and time, the recorded data will display (cid:145)ST(cid:146) or (cid:145)US(cid:146) (stable and unstable) which indicates whether the sample has been stabilized on the pan. This is a useful indicator for the flow conditions of the experiment.

This is a useful indicator for the flow conditions of the experiment. US (unstable) generally indicates the powder is flowing and ST (stable) indicates flow has stopped and it is times to begin or end sampling.Experiments began with the calibration of the balance with an ASTM standard 100g mass.The powder delivery apparatus was positioned over the scale, and powders were deposited into an 8mL narrow-mouth glass laboratory bottle. Twenty samples were taken under each test condition, and mass data from the balance was transferred directly to a PC and stored as a generic data files. The files was then processed in program written in Matlab' for generating plots of the mass accumulation and mass flow rate versus time. To identify time-dependent flow instabilities such as oscillation or arching and their degree of influence on the flow, a program was written for applying a fast Fourier transform (FFT) to the sample data.

The Fourier transform, a mathematical tool frequently used for electrical signal processing, decomposes a complex signal into its basic parts, and display the strength of the revealed signals in a particular range of frequencies. By using an FFT, mass data sampled in discrete time intervals were converted into mass data as a function of frequency. An additional program was also created to filter noisy signals due to possible equipment error.Powder Delivery ApparatusAn extended column for powder above the pipette tip is needed, and additional height is achieved with 30mm long, 8 O.D., 5 I.D Pyrex glass tubing. The tube is suspended vertically by a standard laboratory support stand and clamps.

The tube is suspended vertically by a standard laboratory support stand and clamps. The end of the Pyrex glass tube was ground down, tapering to 7mm so that pipette tips could be conveniently attached and removed during experiments.Modifications to the conditions at the upper free surface of the powder column allow foreach of the mass flow and nozzle test conditions: gravity, closed- top, and pressure- assisted flow. Gravity flow experiments are conducted at standard atmospheric pressure (14.696 psi(101kPa)), which means the top of the powder column is open to air and may interact with the powder particles at the free surface as well as through the nozzle. Closed-top experiments are sealed at the top with a rubber stopper, permitting air to be exchanged for powder particles only through the nozzle.623Figure 4.

Closed-top experiments are sealed at the top with a rubber stopper, permitting air to be exchanged for powder particles only through the nozzle.623Figure 4. Test apparatus (pressure regulators not shown).Pressure assisted experiments are designed to achieve continuous flow in pulsingpowder/nozzle combinations and move cohesive powders by counteracting the viscous drag forces at the orifice. A 250 psi compressed air supply, regulated by two Bellofram' Type 70 precision air regulators, is fed through 1/4(cid:148) O.D., 3/16(cid:148) I.D. nylon tubing to the powder column. Both air regulators are capable of reducing the 250 psi line supply, however the 0-2 psi regulator has a much larger number of turns than the 0-30psi., providing fine tuning in the low pressure range. Two Omega' digital pressure gauges with 0.01 psi resolution are joined to the regulators for adjustment readings.

Two Omega' digital pressure gauges with 0.01 psi resolution are joined to the regulators for adjustment readings. Pressures of 0.25 (1.7kPa), 0.5 (3.4kPa), and 0.75psi (5.2kPa) are chosen for testing with each orifice diameters and particle size. The choice of pressures is based on observing that the minimal regulated pressure of 0.01 psi is raised to 0.25 psi. when added to a full powder column. A stopcock is used above the powder column to stop/start air-presssure assistance to prevent unwanted spraying of particles once the powder column has emptied.Numerical ModelDue to the highly complex multiphase flow behavior inside hoppers, numerical modelsare required for gaining detailed knowledge about gas-solid flow important for designing the proposed powder deposition system. In order to develop a scientific understanding of the flow of powders in our hopper-nozzle designs, numerical models based on kinetic theory of granular multiphase flow will be developed.

In order to develop a scientific understanding of the flow of powders in our hopper-nozzle designs, numerical models based on kinetic theory of granular multiphase flow will be developed. This theory is based on the continuum equations of mass and momentum, and a transport equation for particle fluctuation energy based on the concept of a granular temperature to account for interactions between granular particles, as discussed by Gidaspow [3,4].CFX, a computational fluid dynamics software that has previously been employed withconsiderable success for modeling granular multiphase flow [87-89] will be used to simulate 2-D and 3-D flow models of powders (10nm-100(cid:181)m) inside hoppers (0.1-2.5mm openings). Conservation equations of mass and momentum will be developed for both the gas phase and the particulate phase using the Eulerian approach, and finite volume numerical methods will be employed to solve these equations.

Detailed flow profiles of each phase in the hopper will be simulated, including the velocity and mass flow of particles, local solid fraction, and the pressure distribution of the interstitial gas. The 3-D numerical models will be verified by comparing the624computed total particle mass flow with our experimental results. The effects of powder properties and hopper geometry on the bulk flow behavior of the particles will be studied. Critical conditions leading to undesirable arching, rat-holing and oscillatory flow behaviors will also be investigated.RESULTSFigure 5 shows the accumulated mass v. time, and Figure 6 the corresponding mass flow rate v. time for -140/+170 mesh (90-106(cid:181)m diameter) powder under standard gravity flow from a pipette tip with 1.0mm orifice. This plot is a sample of a (cid:147)continuous(cid:148) flow condition, where a stable (within 0.01g/s) flow can be observed over the full length of the powder column.

Figures 7 & 8.In Figures 7& 8 a (cid:147)continuous(cid:148) mass flow can be achieved with 0.25psi gas assistance for 38- 68(cid:181)m particles and tip diameters down to 0.75mm. The initial mass flow rate achieved is very similar to that of the powder under gravity flow. However, this continuous flow behavior exists for a characteristic powder column length greater than 15-20cm. This phenomenon can be observed by the slight upward turn in mass accumulation (Figure 7), the more obvious increase in mass flow rate beginning after 50 seconds (Figure 8), and significant increase to 0.15g/s almost 3 times as the powder column empties. This is one demonstration of how fine particles that do not flow under standard gravity conditions may be assisted by a very low pressure to achieve comparable mass flow rates.625120120Plot of 10-25 Micron Particles with 1.5 mm Nozzle under 0.75psi Flow, Sample # 1316Plot of Mass Flow rate vs.

Pipet Nozzle Diameter for +140/-170mesh (90-106m icron) Glass Beads14) g (120.5s s a M100.4580.460123456) s / g ( e t a r0.350.310w ol0.25) s / g ( e t a r5F s s a M0.20.15w o l f0.1s s a M00.05012Time (s) 3456011.21.41.61.82Figure 9 &10.Figure 11.Pipet Diameter(m m)However, Figures 9&10 indicate particles as small as 10-38(cid:181)m are very difficult to deliver with a predictable mass flow rate. A higher level of pressure assistance (0.75psi is shown) pushes particles down the powder column at a high rate, yet these very cohesive powders are not being fluidized. By definition, the inter-particle forces of a cohesive powder resist fluidization, and interstitial gas has a harder time permeating the powder column. The flow may be subject to frequent stopping as shown.

The flow may be subject to frequent stopping as shown. For now, highly cohesive powders flow out of the powder column at rates which are too high and unpredictable for incorporation with SLS.Mass Flow Rates v. DiameterUnder gravity flow in pipets, for fixed particle diameter, continuous flow can be observed over a range of orifice diameters. However the flow rate exhibits a nonlinear relationship with respect to orifice diameter (Figure 11). There appears to be optimal combinations of hopper- nozzle diameter to particle diameter. This is helpful for recognizing that there is no simple assumption of a linear relationship between mass flow rates and reduced orifice diameters.

This is helpful for recognizing that there is no simple assumption of a linear relationship between mass flow rates and reduced orifice diameters. Experiments are being developed to determine how this phenomenon relates to the taper angle of the hopper or electrostatic charging at the powder column walls.CONCLUSIONSA preliminary study of the flow of fine powders from small-scale hoppers demonstrates that highly spherical particle 38-100(cid:181)m in diameter may be delivered at continuous mass flow rates with low gas-pressure assistance as small as 0.25psi. Alternate techniques must be developed for delivering very cohesive particles in the range of 0.1-38(cid:181)m. This results of this study are useful for development of a hopper-nozzle array for incorporating multiple powders in the SLS process. Future work will include comparing the results of the numerical model with experimental data to enable further refinement of the model.

PERIMENTAL AND THEORETICAL ANALYSIS OF SCANNING LASER EPITAXY APPLIED TO NICKEL-BASED SUPERALLOYSR. Bansal, R. Acharya, J. J. Gambone, and S. DasWoodruff School of Mechanical Engineering, Georgia Institute of Technology Atlanta, GA 30332-0405AbstractThis paper reports on the experimental development and the theoretical analysis of the scanning laser epitaxy (SLE) process that is currently being investigated and developed at the Georgia Institute of Technology. SLE is a laser-based manufacturing process for deposition of equiaxed, directionally solidified and single-crystal nickel superalloys onto superalloy substrates through the selective melting and re-solidification of superalloy powders. The thermal modeling of the system, done in a commercial CFD software package, simulates a heat source moving over a powder bed and considers the approximate change in the property values for consolidating CMSX-4 nickel superalloy powder.

The theoretical melt depth is obtained from the melting temperature criteria and the resulting plots are presented alongside matching experimental micrographs obtained through cross-sectional metallography. The influence of the processing parameters on the microstructural evolution, as evidenced through observations made from the micrographs, is discussed. This work is sponsored by the Office of Naval Research, through grants N00173-07-1-G031 and N00014-10-1-0526.IntroductionIn order to increase the operating temperature of gas turbine aero engines, investment cast equiaxed turbine component parts are now often replaced with directionally solidified components that have either columnar or single crystal (SX) morphology. During the operation of these SX turbine blades, damage and wear at their leading tip occurs which limits the lifetime of the SX components and necessitates the replacement of the numerous airfoils within the engines.

Due to the high cost of producing SX cast turbine blades, the cost of replacing each blade is several thousand dollars. With each engine containing several hundred blades, the cost to replace every one can reach hundreds of thousands of dollars. Hence, it is of great interest to develop a process which will restore the single crystal microstructure at the damaged location and allow for the blades to be reused rather than scrapped and replaced.Such repairs were first attempted using processes such as Tungsten Inert Gas Welding and Laser melting using a defocused laser (United Technologies Corporation). However, these processes either could not create single crystal deposits, could not be applied to complex geometries, or suffered from crack formation[1]. Laser Engineered Net Shaping, Laser Cladding and Epitaxial Laser Metal Forming have more recently attempted to restore single crystal airfoils to working condition[2].

Laser Engineered Net Shaping, Laser Cladding and Epitaxial Laser Metal Forming have more recently attempted to restore single crystal airfoils to working condition[2]. However, crack formation, columnar to equiaxial transition (CET), oriented to misoriented transition (OMT) and stray grain formation hinders the applicability of all of the above processes to this particular tip repair process [1, 3-5]. This work focuses on thefurther development of a new process, scanning laser epitaxy (SLE), by which arbitrary SX high value components can be repaired.SLE is a laser-based manufacturing process for deposition of equiaxed, directionally solidified and single-crystal nickel superalloys onto superalloy substrates through the selective melting and re-solidification of superalloy powders.

In this particular implementation of the process, a laser beam guided by a set of high-speed galvanometer scanners allows for tight control of the amount of energy being applied to the top of the powder bed as well as the speed at which the melt pool moves across the substrate. Under the proper operating conditions and with enough substrate re-melt, the solidification of the powder melt pool will follow the morphology of the underlying substrate, allowing for directional and even SX growth.Current work focuses on experimentation used to establish an optimal operating range for the various process parameters in the SLE blade repair process. The effects of these different processing parameters on build quality are also examined from the microstructural point of view. In particular, the change of grain structure and morphology along with any angular misorientation is studied in detail.

In particular, the change of grain structure and morphology along with any angular misorientation is studied in detail. A thermal model has also been incorporated in the analysis to predict the melt depth considering the approximate property changes in the powder bed during melting. The model simulates the entire laser scan path and gives some fundamental idea about the effect of repeat scans on preheating, the approximate melt depth and the temperature distribution in the powder bed and substrate.ExperimentalDue to its frequent use in modern SX airfoil designs, the nickel-based superalloy CMSX- 4 was chosen as the material to be used in this evaluation of the SLE process. The CMSX-4 powder, listed in Table 1, was produced by Praxair Surface Technologies using an atomization process and had a particle size ranging from 85-150µm. Initially, the process was tested on rectangular SX cast CMSX-4 coupons having dimensions of 35.56mm x 10.16mm x 2.54mm.

Initially, the process was tested on rectangular SX cast CMSX-4 coupons having dimensions of 35.56mm x 10.16mm x 2.54mm. Each substrate was placed into a 35.56mm x 10.16mm recession cut into an Inconel 625 base plate.CMSX-4Cr 7.6Table 1. Chemical Compositions (wt%) of CMSX-4 W 2.0Co 9.3Mo 0.4Re 1.0Al 12.6Ti 1.3Ta 2.2Hf 0.03During preliminary testing the powder was deposited in various thicknesses, between 1 and 2mm, to find the ideal height of deposit. The powder was held in place above the substrate using wells cut into a Cotronics foam ceramic board that is resistant to thermal shock.After the samples were prepared, they were placed into a Terra Universal atmospheric glove box which was then purged with argon. A 1064nm 2kW Nd:YAG laser beam was used in conjunction with a Cambridge Technologies galvanometer scanner to focus the beam to a diameter of 1.5mm at the top of the substrate.

A 1064nm 2kW Nd:YAG laser beam was used in conjunction with a Cambridge Technologies galvanometer scanner to focus the beam to a diameter of 1.5mm at the top of the substrate. A consistent raster scan pattern across the width of the sample was used to propagate a linear melt pool across each substrate. Three variables were altered between each experimental run: laser power, scan speed, and the number of repeatedNi Balscans at the start of the raster pattern. Table 2 lists the operating ranges for each of the parameters.Table 2. Process parameter ranges found to produce well formed CMSX-4 deposits.Process Parameter Range Powder Thickness Laser Power Scan Speed Initial Repeat Scans1 – 2 mm 400 – 600 W 150 – 250 mm/sec 50 – 150The laser power and speed of the scan were both used to control the amount of energy being applied to the sample during the raster scan pattern.

The number of initial repeat scans affected the formation of a melt pool at the start of the substrate as well as preheat temperature of the entire substrate. Sufficient preheat was required in order to form a proper fusion bond across the whole substrate. Without sufficient preheating, the melt pool would not wet the surface of the underlying substrate and voids were formed along the interface.Microstructural InvestigationEach sample was sectioned along the length and width to allow for views of the microstructure. Using a Buehler automated saw, each sample was first cut lengthwise, and then sectioned by multiple widthwise cuts, as shown in Figure 1.Figure 1. Sections used in metallographic analysisEach section was mounted in Bakelite and ground to a smooth finish; starting at 80 grit paper and progressively increasing the grit to 1200. The samples then underwent a rough polishing operation with 5µm and 3µm diamond solutions.

The samples then underwent a rough polishing operation with 5µm and 3µm diamond solutions. Finally, the samples are smoothed to a mirror finish using a 0.5µm silica suspension. The polished samples were then exposed to Kalling’s No. 2 etchant composed of 50ml Ethanol, 50ml HCl and 2.5g CuCl2 in order to expose the microstructure for microscopy. Imaging was then completed using a Leica DM6000 Optical Microscope at 100x magnification.As demonstrated by the sample in Figure 2, many sections showed a full metallographic bond along the entire length of the sample. The transition from the cast CMSX substrate to the laser processed material is indicated by the sharp change in dendrite size. Also of note are the fully dense porosity free deposit, and the lack of any hot tearing or stress cracking. These traits are consistently found in samples manufactured using the SLE process and are necessary for their implementation on any commercial engine.Figure 2.

These traits are consistently found in samples manufactured using the SLE process and are necessary for their implementation on any commercial engine.Figure 2. Representative lengthwise section of the first half of a CMSX-4 sample with the starting edge on the left side of the imageFigure 3 provides a detailed view of the transition from the coarse dendrites of the cast CMSX substrate to the finer dendrite size created during the experiments. The dendrite arm spacing of the added material is found to be ~15 times finer than in the substrate due to a focused heat input and the increased speed at which the SLE process occurs resulting in a very high temperature gradient as compared to the casting process. The orientation of the columnar growth in the newly deposited material also shows little to no angular misorientation as compared to the underlying substrate.Figure 3.

The orientation of the columnar growth in the newly deposited material also shows little to no angular misorientation as compared to the underlying substrate.Figure 3. Coarse to fine grain transition marking the transition from substrate to depositFigure 4 provides a closer look at the transition from single crystal growth with [001] directionality to a shift in orientation. This change is known as oriented to misoriented transition (OMT) and is caused by a change in direction of the temperature gradient as the substrate becomes increasingly hot toward the end of the scan.Figure 4. Oriented to misoriented transitionOn some samples, the top of the deposit suffered from an area of equiaxial grain growth. This columnar to equiaxial transition (CET) is due to the increase in isotherm velocity near the surface and a decrease in the temperature gradient which causes formation of a constitutionally undercooled zone [1].

Despite these regions of misoriented and equiaxial grain growth, most samples showed an SX deposit height of at least 500µm. In production parts, the misoriented and equiaxial dendrite formations would be removed via a grinding operation or via a re-melt operation if multiple layers were to be built.ModelingComputer simulations were conducted to enable further understanding of the heating and cooling process that leads to the microstructure found in the samples. Initial efforts focused on a thermal model to predict the melt depth along the length of the sample.A multi-domain model with temperature-dependent property values was used to simulate the thermal aspect of the model and to predict the melt depth. The mesh for the given problem was generated in the ANSYS Mesher. The element count for the given case was around 110k with all hexagonal elements and skew not exceeding 0.44.

The mesh for the given problem was generated in the ANSYS Mesher. The element count for the given case was around 110k with all hexagonal elements and skew not exceeding 0.44. The problem was solved in the finite volume based CFD solver, ANSYS CFX, without incorporating the fluid convection.The equation solved for the given model is shown in Equation 1,where, = Density, H = Enthalpy, K = Thermal conductivity, S0 = Heat source. A convection boundary condition was applied along the entire external surface to model the convection heat loss with a high convection coefficient at the top surface (100 W/m2K).The powder bed properties were dynamically modified with the local temperature value as listed in Table 3.Eq. 1Table 3. Powder bed properties for CMSX-4T (K) 300 1636(kg/m3)4350.59 8700K (W/mK) 14.2387 28.4512Cp (J/kgK) 929.94 815.88The moving heat source was modeled with the total input laser power being confined within a 1.5mm beam diameter Gaussian distribution.

The laser initially scans the starting edge for a specified number of iterations, listed as the number of repeat scans, before starting the raster scan over the remainder of the substrate. The moving heat source was simulated following the same path in CFX. The heat source value was also varied to approximate the change in absorptivity for CMSX-4 powder in the final sintered stage [6].The problem was modeled in CFX using the multi-domain approach, with conservation of H- energy being maintained at the interface of the substrate and powder bed. The model did not consider the latent heat since phase change was not yet incorporated.The temperature distribution after the initial repeat scans for one particular case is shown in Figure 5. The bulk preheating obtained, as calculated from the volume averaged quantity, is 47 K for the given case.Figure 5.

The bulk preheating obtained, as calculated from the volume averaged quantity, is 47 K for the given case.Figure 5. Simulated temperature distribution after completing the repeat starting edge scansThe melt depth was computed via tracking of the maximum temperature value obtained at any given plane and comparing it with the liquidus temperature for CMSX-4. The simulated melt depth was then compared with the experimentally obtained melt depth and a reasonable match was obtained over most of the scan length, as shown in Figure 6.Figure 6. Plot of the simulated melt depth, indicated by the red line, superimposed on top of an experimental micrograph run at identical processing parametersInaccuracy in the model may arise due to not incorporating the phase change, the radiation loss from the boundary and approximate convection coefficient used in the model.

Misalignment of the scan path adopted in the model as compared to the actual experimental scan path may also play a role in the relatively larger inaccuracy in the predicted melt depth at the starting edge of the sample. Future simulation work will involve modeling of the microstructure formed from the melt fraction and temperature gradient data.ConclusionsThe intention of the current work was to establish an operating range of input parameters which will create a favorable temperature gradient and solidification velocity that promotes SX morphology. This was complemented by developing a thermal model which could successfully predict the melt depth and to estimate the level of preheating obtained due to starting edge repeat scans.The current effort also leads to some important microstructural observations. The resultant microstructure was substantially finer compared to the investment cast substrate.

The resultant microstructure was substantially finer compared to the investment cast substrate. The deposits did not show cracking or hot tearing once the proper operating range of process parameters was established. The low-angle misorientation did not exceed the allowable range, as evidenced by the microstructural analysis. The height of columnar grain growth obtained was restricted by a flip in grain growth direction near the top trailing edge of the deposit (OMT) as well as due to the formation of an equiaxed cap (CET).

Due to the large height of the SX portion of the deposits produced via SLE, the misoriented and equiaxed portions at the top of the deposit can be removed using a post processing operation while still leaving a significant amount of newly deposited SX material.Future work will encompass the reduction of these misoriented and equiaxed regions near the top of the deposits by utilizing feedback control in order to better control the temperature gradients present during processing. The feedback control will be based upon themodeling of microstructural formation from the simulated melt fraction and temperature gradient data.AcknowledgementsThe authors of this paper would like to thank the Office of Naval Research, and in particular Program Managers Dr. Khershed Cooper and Dr. Ralph Wachter, for their support under grants N00173-07-1-G031, N00014-10-1-0526, N00014-10-1-0756, and N00014-11-1-0670.References1.

perimental assisted design development for a 3D reticulate octahedral cellular structure using additive manufacturingLi YangDepartment of Industrial Engineering, University of Louisville, Louisville, KY 40292Abstract Traditionally it has been difficult to develop and verify designs for 3D reticulate cellular structure. Additive manufacturing provided a feasible alternative for this challenge. In this work, a 3D octahedral cellular structure was designed and investigated. Using a combined method of simulation and experimentation, the mechanical properties of the structures were evaluated. It was found that the cellular structure exhibits unusual size effect that is highly predictable by simulation and experimentation.

It was found that the cellular structure exhibits unusual size effect that is highly predictable by simulation and experimentation. This work established the design-property mapping for the octahedral cellular structure for further design development, and demonstrated the feasibility of applying this type of structure in sandwich panel applications.Introduction Cellular structures have been drawing increasing attentions in recent years due to their potentials in high performance-to-weight ratio applications. Sometimes referred to as “metamaterials”, cellular structures could achieve tailored physical and mechanical properties via the control of interior geometries [1, 2]. Additive manufacturing (AM) has been increasingly often used to manufacture cellular structures due to its freeform capability with complex geometries [3-6]. On the other hand, the design theory for cellular structures is relatively underdeveloped, especially for the cellular structures with 3D architectures.

On the other hand, the design theory for cellular structures is relatively underdeveloped, especially for the cellular structures with 3D architectures. By far only several of such 3D cellular structures have been investigated in details for design development, including octet-truss structures [7-10] and re-entrant auxetic structures [4, 11-14]. It has been suggested that the development of design tools for 3D cellular structures could be facilitated by AM technologies [15]. With the fabrication capability of AM, cellular designs could be verified by experimentation, and other non-geometrical design factors such as manufacturing factors and boundary factors could be conveniently identified and investigated. In this paper, works were focused on the preliminary modeling of an octahedron 3D cellular structures with electron beam melting (EBM) process.

In this paper, works were focused on the preliminary modeling of an octahedron 3D cellular structures with electron beam melting (EBM) process. The octahedral structure was of interest to this study due to its relatively straightforward deformation mechanism and the potentials in high energy absorption sandwich structure applications. The approach of this study took the same methodology used previously in the development of auxetic structure with EBM [4, 11-14], in which analytical modeling, finite element analysis and experimentation were used in combination in order to gain insights into the design issue.Modeling of structure The octahedral structure is a special type of 3D cellular structures. Unlike many other commonly seen cellular structures, this octahedral structure does not have tetrahedral or triangular sub-structures. As shown in Fig.1(b), the octahedral structure can be represented by an octahedral volume with eight sloped struts.

As shown in Fig.1(b), the octahedral structure can be represented by an octahedral volume with eight sloped struts. This structure has been previously studied experimentally, although the work was focused on process perspectives [16] The Maxwell stability criterion number for the octahedral structure is M<0 [17], which indicates that the structure does not have the same kinematic stability as compared to some of the other 3D cellular structures such as octet-truss. It’s worth noting that for structural analysis purpose, the unit cell of the octahedral cellular structure could be constructed around a joint. As shown in Fig.1(c), the unit cell structure possesses a central joint with eight struts radiate out. From the construction of the unit cell, it is obvious that the octahedral structure has the same geometrical patterns in all the three principal directions, and therefore would exhibit the same loading response mechanism in all the three principal directions.

As a result, for the purpose of modeling, only one principal direction needs to be investigated.(a) lattice structure(b) Unit cell construction 1 Fig.1 Octahedral cellular structure(c) Unit cell construction 2In this study, work was focused on the uniaxial properties of the structure. Consider a remote compressive stress σ applied on the structure along the z direction. The loading of the unit cell structure is shown in Fig.2(a). From structural symmetry, it is apparent that all the struts are under the same loading conditions. Fig.2(b) further illustrates the loading condition of an arbitrary strut.

From structural symmetry, it is apparent that all the struts are under the same loading conditions. Fig.2(b) further illustrates the loading condition of an arbitrary strut. From Fig.2, the following force components could be obtained:F z1 4L L 1 2 z(1)M11 2F L 3 zsin1 8   zzL L L 1 2 3sin z(2)where L1, L2, are the dimensions of the unit cell in x and y directions, and L3 is the length of a half-strut, and θz is the slope angle of the strut in relation to the z axis: zsin1 2 L 1 2 L 32 L 2(3)(a) Loading of unit cell(b) Analysis of individual strutFig.2 Octahedral unit cell under compressionUnder the loading condition shown in Fig.2(b), the strut undergoes a bending/shearing combined deformation. It has been shown before that the deformation along the strut axis is insignificant for metal cellular structures under non-catastrophic stress levels, therefore is not considered [18].

It has been shown before that the deformation along the strut axis is insignificant for metal cellular structures under non-catastrophic stress levels, therefore is not considered [18]. Apply Timoshenko beam analysis, the deformation of the strut in the z direction and the direction perpendicular to the z direction (not labeled in Fig.2(b)) can be obtained as: z L 3sin z  M L 1 3 EI 66F sin z GA 5 z  (3)z   L 3cos z  M L 1 3 EI 66F sin z GA 5 z  (4)where E, G are the modulus of elasticity and shear modulus of the solid material, I is the second moment of inertia of bending in the plane shown in Fig.2(b), and A is the cross sectional area.

If the deflection Δz┴ is further decomposed into deformations in the x and y directions, Δx and Δy, then: x L 3coscos   zz  M L 1 3 EI 66F sin z GA 5 z  (5) y L 3sincos   zz  M L 1 3 EI 66F sin z GA 5 z  (6)Therefore, the modulus of the unit cell can be expressed as:EzL 3cos  z z z H 2 z z z 2L L L 1 2 3sin2H    z2 L 3 EI 483 GA 10  (7)WherezH2L 3cos22 L 31 4(2 L 12 L 2), as is also shown in Fig.2(a).The strength of the structure could be roughly estimated as the onset of the yield in arbitrary strut. This could provide an estimation of yield strength of the structure. Since it is known that thin features produced by EBM process exhibit significant surface defects [18‐21], it is reasonable to use more conservative estimations for the strength prediction of cellular structures.

The normal and shear stress of the strut shown in Fig.2(b) can be determined from the bending moment, normal force and shear force as:  1M u F z1 Icos A z zL L L 1 2 3 I 8sin zu zL L 1 2 4cos A z(8)  1sinz F Ib zD(9)Where u is the distance from the geometrical center of the cross section to the location of interest, D is the moment of area of the cross section, and b is the width of the cross section at the location of interest. Apply Von Mises Criterion, the maximum allowable stress level σm that results in the onset of structure yield is: mL L 1 22 L 32sin 2 I 2 zu242 cos 2 A2 z26sin z 2 2 I bD2 Y(10)Fig.3 and Fig.4 show the modulus and yield strength of the octahedral structures as functions of geometrical design parameters.

The mappings were generated to illustrate the range of mechanical properties achievable by geometrical designs, and it is obvious from the modeling that the mechanical properties of the octahedral structures could be tailored at a wide range. The deformation of the octahedral structure is largely dominated by bending and shearing, therefore it is a bending dominated cellular structure [17]. The complete lack of tension‐compression strut in the unit cell also determines that bending/shearing is the only active deformation mechanism for the octahedral cellular structure. It was predicted that bending dominated structures could dissipate considerable amount of energy upon plastic deformations, therefore, the octahedral structure might possess good potential in energy absorption applications.Fig.3 Modulus of the octahedral structureFig.4 Yield strength of the octahedral structureAnother interesting characteristic of the octahedral structure also originates from its geometry.

Throughout the modeling process, no assumption of infinite structure was required, which is often adopted for the other cellular structures in order to simplify the modeling. As a result, the equations derived for the octahedral structure is expected to be applicable for even lattices with small numbers of unit cell repetitions. This is in significant contrast with the general theory for cellular structures, in which about 8‐10 unit cells are required in order for the structures to minimize the boundary effects [22, 23]. This property could also be very useful in applications, especially for the design of sandwich structures with very limited spaces. For these structures, only one or two layers of periodic cellular cores are oftenallowed, and the ability to accurately predict the mechanical properties of the core structure at this scale is highly desired.In order to verify this observation, finite element analysis (FEA) was performed with octahedral structures of various sizes.

Models with unit cell numbers (x x y x z) of 1x1x1, 2x2x1, 4x4x1, 8x8x1, 1x1x2, 2x2x2, 4x4x2, 8x8x2, 1x1x4, 2x2x4, 4x4x4, 8x8x4 and 8x8x8 were constructed and analyzed in SolidWorks using COMSOL simulation module. The numbers of unit cell repetition was chosen for the easiness of modeling. The simulations were performed with an office workstation, and it was noted that the computational load of the simulation quickly exceeded the practical level when the total number of unit cell exceeded the 512 level. Therefore, no simulations beyond 8x8x8 were performed. An arbitrary unit cell design was used for the simulations with L1=20mm, L2=18.74mm, H=20mm and square strut of t=1mm. While the type of material is not of concern of the simulation, the material was defined as Al‐ 6061‐T6 alloy. In the simulation, a curvature based mesh with mesh size of 0.9mm was used throughout the simulation.

In the simulation, a curvature based mesh with mesh size of 0.9mm was used throughout the simulation. The mesh size was tested by performing static compression simulations with different mesh sizes, and it was found that the results of the relatively coarse 0.9mm mesh size were adequately accurate for the purpose of this study. Fig.5 shows the boundary condition setup for the simulation. The structure was bonded to the two platens on both sides of the loading direction, and the stiffness of the platens was set to be significantly higher than the structure material in order to minimize the analysis error. One platen was fixed, and the other was loaded with uniaxial compressive stress. Since elastic analysis was performed, only the size effect of elastic modulus was verified.Fig.5 Boundary condition setup for the FEA simulation of size effectThe results of the simulation are listed in Table 1.

It was found that despite the prediction, the octahedral structure exhibits significant size effect. However, upon close look, the octahedral structures exhibit a counterintuitive size effect in the direction of the loading. While the modulus of the structure increases with increased number of unit cells parallel to the loading direction, it exhibits a steady reducing trend as the number of unit cells in the loading direction (z direction) increases, which is further illustrated in Fig. 6. For regular cellular structures, size effect tends to soften the structure due to the loss of symmetry at the boundaries. However, for octahedral structure, there exists an opposite trend. With fewer numbers of layers, the structure actually becomes stiffer. It was also observed that the size effects in both cases follow a very predictable logarithm trend.

With fewer numbers of layers, the structure actually becomes stiffer. It was also observed that the size effects in both cases follow a very predictable logarithm trend. This unconventional phenomenon was mentioned in previous studies with similar structures, although no further details were discussed [25].Design1x1x1 1x1x2 1x1x4 2x2x1 2x2x2 2x2x4 4x4x1 4x4x2 4x4x4 8x8x1 8x8x2 8x8x4 8x8x8FEA modulus (MPa) 2.751 2.721 2.702 42.955 3.992 3.256 93.067 51.394 6.024 125.786 96.177 53.121 8.060Calculated modulus (MPa) 2.300 2.300 2.300 2.300 2.300 2.300 2.300 2.300 2.300 2.300 2.300 2.300 2.300FEA modulus – Free boundary (MPa) 2.563 2.623 2.654 2.620 2.685 2.714 2.656 2.719 2.742 ‐ ‐ ‐ ‐Table 1 Size effect evaluation for octahedral structure(a) Lateral size effect(b) Layer‐wise size effectFig.6 Size effect of the octahedral structure modulus by FEAFurther investigation revealed the cause of this unconventional size effect.

As shown in Fig.7, when the number of layers increases, the octahedral structure starts to exhibit specific diagonal pattern of stressconcentration. On the other hand, when the boundary conditions were adjusted to allow for relative sliding between the compressive platens and the octahedral structures, this size effect largely disappeared, as shown in the last column of Table 1. This phenomenon is likely originated from the boundary restrictions. As is commonly known, a solid exhibits upsetting when compressed between two platens, which is caused by the friction between the solid and the platens [24]. The upsetting phenomenon does not appear to be reported in any previous studies, and it could be intriguing consider the fact that cellular structures are largely dis‐continuous solids in 3D space, therefore the stress distribution would be largely geometry dependent.

It was expected that this boundary effect would be difficult to eliminate in practice since for most applications the cellular structures would be restricted at the boundaries. However, due to the difficulty of modeling upsetting phenomenon, further works exceeded the scope of this paper, and will be addressed in future studies. On the other hand, the good predictability of this size effect as shown in Fig.6 suggested that a relatively effective design method could be developed based on a reasonable set of experiments.Fig.7 Stress concentration of octahedral structureExperiments and discussions A random parameter set was chosen to verify the accuracy of the modeling for the octahedral structure. The parameters and the theoretical predictions are shown in Table 2. The structures were modeled in SolidWorks and fabricated in the Arcam EBM S400 machine using Ti6Al4V‐ELI as material. In the CAD models, the cross sections of the struts were designed to be square.

The structures were modeled in SolidWorks and fabricated in the Arcam EBM S400 machine using Ti6Al4V‐ELI as material. In the CAD models, the cross sections of the struts were designed to be square. However, due to the feature resolution limit of the EBM process, the actual strut cross sections of the samples exhibit irregular shapes, therefore a round cross section approximation was adopted in the calculation of models. Octahedral structures with 1x1x1, 1x1x2, 1x1x4, 2x2x1, 2x2x2, 2x2x4, 4x4x1, 4x4x2, 4x4x4 numbers of unit cells were designed, and the samples were fabricated in three builds, with one sample for every design fabricated in each build.DesignL1 (mm) 10t (mm) L2 (mm) 1 8 Table 2 Design parameters for the octahedral samplesH (mm) 10Theo. Modulus (MPa) Theo. Strength (MPa)44.7731.935Fig.8(a) shows the fabricated octahedral lattice samples. The overall dimensions of each fabricated sample were measured by caliper.

Modulus (MPa) Theo. Strength (MPa)44.7731.935Fig.8(a) shows the fabricated octahedral lattice samples. The overall dimensions of each fabricated sample were measured by caliper. Compressive test was performed for the samples with an Instron 5569A universal tester. The samples were fixed between two steel platens, 50000N load cell was used for samples with 4 unit cell lateral repetitions, and 5000N load cell was used for the other samples. The displacement of the crosshead was used as the total deformation of the structures consider the small modulus of tested samples. Fig.8(b) shows the actual setup of the compressive testing for the octahedral lattices. The testing results are shown in Table 3. The dimensions and mechanical properties of the samples were quite consistent, indicating a relatively stable manufacturing quality at strut size of 1mm.

It was assumed that the mechanical properties of the Ti6Al4V material fabricated by EBM is isotropic, therefore the results obtained from experimentation could be directly compared with the prediction.(a) Octahedral lattice(b) Compressive testFig.8 Octahedral lattices for experimental verificationDesign1x1x1 1x1x2 1x1x4 2x2x1 2x2x2 2x2x4 4x4x1 4x4x2 4x4x4Modulus (MPa) 27.27±0.34 9.68±0.06 9.94±0.04 7.86±0.01 19.88±0.18 28.76±0.55 9.92±0.04 7.90±0.07 40.08±0.20 31.41±1.14 7.92±0.05 9.90±0.06 33.46±1.82 9.72±0.16 19.82±0.01 15.80±0.03 19.80±0.03 15.79±0.02 19.73±0.11 34.61±1.69 19.94±0.06 15.75±0.03 40.18±0.19 37.96±0.78 31.67±0.05 39.58±0.06 39.42±3.79 31.60±0.03 39.54±0.05 19.83±0.13 37.41±1.50 31.66±0.03 39.61±0.07 40.18±0.18 39.11±1.16 Table 3 Compressive testing results for octahedral latticeD1 (mm)D2 (mm)D3 (mm)9.89±0.21Strength (MPa) 1.53±0.01 1.63±0.08 1.11±0.11 2.24±0.05 1.93±0.04 2.00±0.01 3.10±0.37 2.32±0.03 2.39±0.08Due to the difficulty of sample orientation with the large aspect ratio designs (1x1x2 and 1x1x4) during compressive testing, the results for these samples were expected to have relatively large errors.

Inaddition, it was also observed that the samples of 1x1x4 design exhibited macroscopic elastic buckling during the compression test, which explained the considerably lower strength of the 1x1x4 design compared to the other designs. Overall, the modulus and strength values of the samples are close to the theoretical predictions, and no significant size effect was observed. Upon close inspection, it was found that the compressive test platens were indented and scratched during the compressive testing, which indicated that under the combination of shearing and compressive stress, there existed relative motions between the samples and the platen surfaces.

Therefore, although the octahedral lattice samples still exhibited stress concentration patterns as shown in Fig.9, the boundary constrains were not ideal due to the relative motion between the samples and the platens, which resulted in mechanical properties that are somewhat close to the theoretical prediction.(a) 4x4x4 lattice(b) 2x2x4 latticeFig.9 Stress concentration pattern of octahedral lattice during compressionConsider the fact that in actual applications the octahedral structures are likely to be fully constrained, the relative motions between the cellular structures and the surrounding structures will not be acceptable. Therefore, a second set of experiments were performed with redesigned samples. In order to ensure full constraint of the octahedral lattices, sandwich panel structures were designed. For the core of the sandwich structures, the same octahedral designs were adopted. The skin thickness of the sandwich structures were 1mm for all the designs.

For the core of the sandwich structures, the same octahedral designs were adopted. The skin thickness of the sandwich structures were 1mm for all the designs. The designs were fabricated with the same process parameters and build setups. Since the skins were also thin features, all the samples were fabricated using the default network process theme provided by the machine. The sandwich models were rotated so that the skins aligned in the vertical orientations, therefore eliminated the potential issue with support requirements. The fabricated samples were measured and tested in the same manner. Fig.10 shows the fabricated sandwich samples and the setup for the compressive test.(a) Fabricated sandwich samples(b) Compressive testFig.10 Octahedral sandwiches for experimental verificationThe results of the experiment are listed in Table 4. As could be observed clearly, the octahedral sandwich structures exhibited significant size effect.

As could be observed clearly, the octahedral sandwich structures exhibited significant size effect. Fig.11 further illustrates the size effects of elastic modulus for the octahedral sandwiches. The size effect clearly showed the same pattern and logarithm trend as indicated by the predictions (Fig.6), and it was obvious that the size effect becomes apparent when the total number of layers exceeds the lateral unit cell repetition. In fact, from Table 4 it was also apparent that the strength of the octahedral sandwich also exhibits the same size effect, which is plotted in Fig.12 for clarity of information. With more unit cell repetitions in the lateral directions or less layers, the octahedral sandwich becomes stronger in compression. It could also be observed from Fig.11 that when the size effect of the octahedral sandwich is minimized, the modulus values of the sandwich structures are consistently lower than the predicted value (44.773MPa).

This could be partly contributed by the manufacturing defects. It has been reported that cellular structures fabricated by EBM process possess a variety of defects [11, 13, 15, 18], which could significantly affect the mechanical properties of the parts. In addition, due to the build orientations selected, one side of the octahedral core of each sample was in contact with the substrate. Upon the parts removal, the joints of the octahedral lattices on these sides were partly damaged. For samples with large lateral unit cell repetitions, the effect of such defects was reduced.

However, with smaller samples, these defects were more likely to result in the loss of mechanical performance, as was clearly suggested in the experimental results.Design1x1x1 1x1x2 1x1x4 2x2x1 2x2x2 2x2x4 4x4x1 4x4x2 4x4x4Modulus (MPa) 24.23±1.63 7.11±0.13 24.35±2.40 7.25±0.21 7.30±0.23 26.74±3.78 15.56±0.83 19.86±0.04 12.32±0.02 111.23±2.02 32.49±1.18 15.30±0.29 19.85±0.04 22.38±0.03 15.31±0.37 19.89±0.03 42.79±0.06 28.45±1.16 30.88±0.22 39.63±0.12 12.15±0.01 132.28±8.39 31.15±0.28 39.69±0.03 22.38±0.04 125.98±1.18 47.15±1.29 31.09±0.33 39.63±0.06 42.73±0.01D1 (mm)D2 (mm)D3 (mm)9.88±0.00 9.94±0.14 9.97±0.2012.11±0.02 22.29±0.10 42.44±0.43Strength (MPa) 1.39±0.17 1.35±0.02 1.42±0.13 3.79±0.16 1.59±0.06 1.57±0.073.60±0.21 2.72±0.12Table 4 Table 3 Compressive testing results for octahedral sandwich(a) Lateral size effect Fig.11 Size effect of the octahedral sandwich modulus by experimentation(b) Layer‐wise size effect(a) Lateral size effect (b) Layer‐wise size effect Fig.12 Size effect of the octahedral sandwich strength by experimentationConclusions	 The design of cellular structures with AM technologies possesses great potential for many future applications with high performance lightweight structures.

In this study, an octahedral cellular structure was modeled and tested. Using a combination of theoretical modeling, simulation and experimentation, insights were efficiently obtained for this cellular structure. Modeling work successfully established an accurate analytical model for the modulus prediction of the octahedral structure. It was also found that the octahedral cellular structure exhibits a highly predictable size effect, which could be utilized in the design of sandwich structures with thickness restrictions.

It was also found that the octahedral cellular structure exhibits a highly predictable size effect, which could be utilized in the design of sandwich structures with thickness restrictions. The good agreement between the experiment and the theory for the octahedral structure also enables high fidelity design and manufacturing of this type of cellular structures for applications.Acknowledgement	 This study was performed in the Rapid Prototyping Center (RPC) in University of Louisville, and the author would like to acknowledge Tim Gornet, Joe Vicars, Gary Graf of the RPC and Hengfeng Gu for their support and help with this project.Reference [1] Lorna J. Gibson. Cellular solids: structure and properties, 2nd edition. Cambridge University Press, New York, NJ, USA. 1997.[2] M.F. Ashby, A. G. Evans, N. A. Fleck, L. J. Gibson, J. W. Hutchinson, H. N. G. Wadley. Metal Foams: A Design Guide. Butterworth-Heinemann, Boston, US. 2000.[3] O. Cansizoglu, O, Harrysson, D. Cormier, H.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEXPERIMENTAL CALIBRATION OF NANOPARTICLE SINTERING SIMULATIONObehi G Dibua*, Anil Yuksel*, Nilabh K Roy*, Chee S Foong†, Michael Cullinan*Department of Mechanical Engineering, University of Texas at Austin, Austin, TX 78712 †NXP Semiconductors, Austin, TX 78735AbstractMicroscale Selective Laser Sintering models have been built as a basis to predict the properties of sintered nanoparticles under isothermal heating. These models use Phase Field Modelling (PFM) to track the diffusion of nanoparticles, resulting in properties such as the change in relative density and shrinkage of the sintered bed with time. To ensure the accuracy of these PFM models, experimental validation has to be done.

To ensure the accuracy of these PFM models, experimental validation has to be done. This paper presents the experimental procedure and results for isothermally heating nanoparticles up to 450 – 600°C, at sintering times varying from 1 to 45 minutes. Measurement uncertainties are calculated from deviations in calculating the density. Experimental results from this process are then used to calibrate the simulation to determine the number of simulation timesteps which correspond to a minute of physical time. The calibration constant derived is then used to map simulation constants to physical constants. These constants are later compared to bulk properties.IntroductionMicroscale Selective Laser Sintering is a promising process which looks at the creation of sub-micrometer parts through the sintering of nanoparticles [1]. These nanoparticles fuse together using energy absorbed from a laser source.

These nanoparticles fuse together using energy absorbed from a laser source. With the development of this process, there is a concurrent need for models able to predict properties of sintered parts. One such model is presented in [2] and looks at the changes in the rate of densification and shrinkage during sintering. The simulation presented in this model uses Phase Field Modelling to track the diffusion of a large number of particles, 43 particles in a 1 by 1 micrometer bed, and 134 particles in a 2 by 2 micrometer bed. The simulations in this paper look at an isothermally heated bed with simulation parameters which are set constant for all particles in the bed. These simulation parameters are temperature dependent, as such, they need to be calibrated against experimental data, to map simulation constants to corresponding physical temperatures.

These simulation parameters are temperature dependent, as such, they need to be calibrated against experimental data, to map simulation constants to corresponding physical temperatures. Copper nanoparticles were chosen for the initial calibration of these simulation constants.Metal nanoparticles such as silver, gold and copper have been studied in recent years due to the high electrical conductivity of these materials [3]. However, most of the sintering experiments carried out for these particles center around measuring electrical conductivity [4,5]. Few researchers have looked into measuring changes in relative density of sintered nanoparticles. Wang et al, studied the densification of Beryllium oxide nanoparticle powders, finding that these particles began to sinter at temperatures lower than the conventional sintering temperatures [6]. There have been no studies done on the densification of isothermally heated nanoparticle inks.

There have been no studies done on the densification of isothermally heated nanoparticle inks. This study shows densification results for the sintering of copper nanoparticles. The experiments are carried out by changing heating times and temperatures. The relative change in density curve1763is determined from analyzing these experiments, and the curve fit is then calibrated against the relative density curve fit from simulations.ExperimentsSintering experiments were carried out on copper nanoparticle inks from Intrinsiq [7]. The first step of the experimental procedure involved drying out the ink. 2ml of copper nanoparticle ink was dispensed into a glass petri dish using a rubber pipette. The petri dish of ink was dried on a hot plate at around 95˚C for 16 hours. After the solvent in the ink dries off, dried copper flakes of coated nanoparticles are scraped off the petri dish with a flat spatula. These dried flakes are then put into crucibles.

After the solvent in the ink dries off, dried copper flakes of coated nanoparticles are scraped off the petri dish with a flat spatula. These dried flakes are then put into crucibles. Pressure is applied to form the flakes into pellets in the crucibles. These crucibles are put into a furnace and subject to isothermal heating. In the furnace, the coating around the particles dry off and the nanoparticles sinter together into a solid pellet. The flow of the experimental procedure is shown in the images in Figure 1. The sintering experiments were carried out under flowing Argon and Hydrogen to control oxidation of the copper nanoparticles. The ratio of Argon to Hydrogen gas flow is 300:100.Figure 1. Experimental procedure. a. Copper nanoparticle ink. b. Dried ink. c. Scraped off dried flakes. d. Pellets in crucible before sintering. e. Pellets in crucible after sinteringSEM images were taken of the flakes before sintering, and the pellets after sintering.

c. Scraped off dried flakes. d. Pellets in crucible before sintering. e. Pellets in crucible after sinteringSEM images were taken of the flakes before sintering, and the pellets after sintering. These images in Figure 2, show that before sintering the particles are discrete and can be seen separate from each other. After sintering, the images of the pellets show that necks have formed between the nanoparticles after heating.1764Figure 2. SEM Images of sintered nanoparticles. a. Before sintering. b. After sinteringAfter the experiments were done density measurements were taken. The density measurement is calculated as the ratio of mass to volume. The mass was measured using a digital weight scale. The volume was calculated from the measured height of the pellets. The height measurements were taken from the deflection of stoppers. A plastic stopper was put into the empty crucible and the deflection of the stopper was measured.

The height measurements were taken from the deflection of stoppers. A plastic stopper was put into the empty crucible and the deflection of the stopper was measured. The height of the pellet was calculated from the difference between the stopper deflections between the empty crucible and the deflections with the pellet in the crucible. The relative density was calculated from the density measurements using Eqn 1.𝜌𝜌𝑖𝑖 − 𝜌𝜌𝑜𝑜 𝜌𝜌𝑖𝑖𝜌𝜌𝑟𝑟𝑟𝑟𝑟𝑟 = Where ρo is the initial density and ρi is the final density.𝐸𝐸𝐸𝐸𝐸𝐸. 1The sintering experiments were carried out at 450, 500, 550 and 600˚C and the relative density data at these temperatures were fit to an expression of the form in Eqn 2. This form of the equation was modified from [8] where the diffusion coefficient follows an exponential decay fit similar to Eqn. 2.𝐾𝐾2 𝑡𝑡+𝐾𝐾3−+𝐾𝐾4𝜌𝜌𝑟𝑟𝑟𝑟𝑟𝑟 = 𝐾𝐾1𝑒𝑒 + 𝐾𝐾5 𝐸𝐸𝐸𝐸𝐸𝐸. 2 A plot of the experiment data points and the best fit curve for the decay in Eqn. 2 is shown in Figure 3.

2.𝐾𝐾2 𝑡𝑡+𝐾𝐾3−+𝐾𝐾4𝜌𝜌𝑟𝑟𝑟𝑟𝑟𝑟 = 𝐾𝐾1𝑒𝑒 + 𝐾𝐾5 𝐸𝐸𝐸𝐸𝐸𝐸. 2 A plot of the experiment data points and the best fit curve for the decay in Eqn. 2 is shown in Figure 3. A consolidation of all these plots is shown in Figure 4. From Figure 3, it can be seen that sintering is characterized by an initial rate of rapid densification and as the sintering time proceeds the rate of densification approaches a steady state value. Figure 4 shows that as the temperature increases, the amount of time it takes for the relative density to reach steady state decreases. These curves from the experiments were set against the results from the simulation and the appropriate simulation constants and simulation time calibration constant were calculated. The process of deriving this is discussed next.1765Figure 3. Experimental data and curve fit at a. 450˚C. b. 500˚C. c. 550˚C. d.

The process of deriving this is discussed next.1765Figure 3. Experimental data and curve fit at a. 450˚C. b. 500˚C. c. 550˚C. d. 600˚CRelative Density Change wtih Timey t i s n e d n0.40.350.3ie g a n h c e v i t a e Rl0.250.20.150.10.05450C500C550C600C001020Sintering time (minutes)Figure 4. Consolidation of experiment fit plotsSimulation CalibrationThe Phase Field Simulation (PFM) in [2] gives results for the diffusion of nanoparticles in a powder bed. The beds are generated using a bed generation simulation [9] to match actual physical beds. An initial configuration of a powder bed is shown in Figure 5.1766zxyFigure 5. Initial Powder Bed ConfigurationPrevious results in the PFM diffusion simulation show that the process of sintering is characterized by an initial stage of neck formation between particles in contact. After the neck forms, the diffusion process shows neck growth and grain boundary migration.

After the neck forms, the diffusion process shows neck growth and grain boundary migration. The analysis in these simulations were done on a center box of the simulation bed, to make the data collated impervious to edge effects. This process of diffusion in the center of the bed is shown in the images in Figure 6.Figure 6. Densification in the center of the simulation bed after a. 0 timesteps. b. 80000 timesteps. c. 160000 timesteps. d. 480000 timesteps. e. 1200000 timesteps. f. 2400000 timesteps1767The relative density of these beds was calculated similar to the experiments using Eqn. 1. The results from the relative density calculation is shown in Figure 7. The error bounds in this calculation is derived from looking at the deviation in the relative density measurements as a result of changing the location of the analysis box in the center of the simulation.

The trend in the relative density curve in Figure 7 matches the trend seen in the experiment curves where sintering is characterized by a fast rate of densification which slows down to steady state with simulation time.0.850.80.75y t i s n e d e v i t a e Rly t i s n e de v i t a e Rl0.70.650.60.550.500.511.522.5Simulation timeSimulation timesteps10Figure 7. Relative density curve from simulationsThe rate of densification in this simulation are set by the temperature dependent constants. These constants are gradient terms related to surface energy, βρ, and grain boundary energy, βη, coefficients as well as the diffusion constants, surface Dsurf, grain boundary, Dgb, and volume Dvol. These constants were changed to get the closest fit for the simulation density curve to the experiments curve for each temperature. The time calibration constant, A, is defined as in Eqn. 3.𝑡𝑡𝑠𝑠𝑖𝑖𝑠𝑠(𝑡𝑡𝑚𝑚𝑚𝑚𝑒𝑒𝑡𝑡𝑡𝑡𝑒𝑒𝑡𝑡𝑡𝑡) 𝐴𝐴𝑡𝑡𝑟𝑟𝑒𝑒𝑒𝑒(𝑚𝑚𝑚𝑚𝐸𝐸) = 𝐸𝐸𝐸𝐸𝐸𝐸.

The time calibration constant, A, is defined as in Eqn. 3.𝑡𝑡𝑠𝑠𝑖𝑖𝑠𝑠(𝑡𝑡𝑚𝑚𝑚𝑚𝑒𝑒𝑡𝑡𝑡𝑡𝑒𝑒𝑡𝑡𝑡𝑡) 𝐴𝐴𝑡𝑡𝑟𝑟𝑒𝑒𝑒𝑒(𝑚𝑚𝑚𝑚𝐸𝐸) = 𝐸𝐸𝐸𝐸𝐸𝐸. 3 A defines the number of simulation timesteps that corresponds to time in minutes. This constant is determined using a bracketing algorithm that converges to the value of A which minimizes the error between the simulation results and the experiment results. The results fitting the simulation to experiments are shown in Figure 8.1768Figure 8. Comparing experimental fit to simulations for a. 450˚C. b. 500˚C. c. 550˚C.The results for the time calibration values are shown in Table 1. The Table shows the results for the range and averages of the error and time calibration values. The range included in these values, like the error bars in Figure 6, are a result of carrying out the calibration analysis on different boxes in the center of the simulation bed. The data in Table 1 shows an overall average error of about 9% between the simulation and experiments.

The data in Table 1 shows an overall average error of about 9% between the simulation and experiments. The time calibration factor shown in the table has an average of 20% deviation between each temperature.Table 1. Calibration Results: Time and ErrorTime Calibration, A (timesteps/minute)Error (%)450˚C 500˚C 550˚CMinimum Maximum Average Minimum Maximum Average 13.54 7.01 6.76101070 57644 71110388864 277293 383773215319 167598 2501023.22 3.57 3.9930.14 10.91 9.791769Once calculated, the time dependent simulation constants were mapped to constants withactual physical meanings. The diffusion constants were mapped using Eqn. 4.22𝐷𝐷𝑠𝑠𝑖𝑖𝑠𝑠𝐴𝐴 60𝑆𝑆𝑡𝑡𝑚𝑚𝑝𝑝𝑒𝑒𝑝𝑝𝑡𝑡 𝑡𝑡𝑚𝑚𝑚𝑚𝑒𝑒𝑡𝑡𝑡𝑡𝑒𝑒𝑡𝑡𝑡𝑡𝑐𝑐𝑚𝑚 𝑡𝑡𝐷𝐷𝑠𝑠𝑖𝑖𝑠𝑠 � � 𝐸𝐸𝐸𝐸𝐸𝐸. 4 Where Dsim represents the simulation diffusion coefficients and Dactl represents the diffusion coefficients when mapped to the corresponding units in a CGS system. A is the time calibration constant in Eqn.

A is the time calibration constant in Eqn. 3 having units of timesteps/minute and S is the size calibration constant. S here is set as 944822 pixels/cm. The values used for the diffusion constants in the simulation where mapped to physical units and are shown in Table 2. Surface, grain boundary and volume diffusion coefficients in the simulation follow the ratio 1000:100:1 respectively.2 => 𝐷𝐷𝑎𝑎𝑎𝑎𝑡𝑡𝑟𝑟 �� =>Table 2. Calibration Results: Diffusion constantsSurface Diffusion coefficients (cm^2/s) AverageUncertainty450˚C 500˚C 550˚C2E-07 2.2E-07 3E-071E-07 7.9E-08 1E-07The diffusion constants used in the simulations at 500˚C were compared to bulk measurements from literature. As of this point there has been no experimentally determined surface diffusion coefficients for copper nanoparticles to compare against the constants in Table 2. Bonzel and Gjostein [10] found that at 500˚C the surface diffusion coefficient of bulk copper is 1.91E-07 cm2/s.

Bonzel and Gjostein [10] found that at 500˚C the surface diffusion coefficient of bulk copper is 1.91E-07 cm2/s. This value falls within the same order of magnitude as the value obtained from the simulations. Additionally, since the ratio of surface area to volume is higher for nanoparticles, the surface diffusion coefficient for nanoparticles is expected to be higher than that for bulk, which is in agreement with the results presented. This comparison is further validation of the model.Future WorkGiven the simulation work and experiments carried out to calibrate the simulation, the next step is to quantify the error in the simulation. As shown in the uncertainty in changing the position of the analysis box, there is a significant amount of error that comes with the configuration of the particles in the bed.

As shown in the uncertainty in changing the position of the analysis box, there is a significant amount of error that comes with the configuration of the particles in the bed. To completely quantify this uncertainty, different simulation beds have to be tested to determine the changes in calibration constants and rate of densification with changing the initial configuration of the bed. Next, given the map of simulation constants to physical temperature, the simulation would have to be transitioned to include transient heating more representative of a temperature gradient. This temperature gradient is important because in a typical laser sintering process, there is a temperature difference between the particles in the laser irradiated spot and the surrounding particles outside the laser heating zone. Also, the simulation has to be transitioned to include multiple layers.

Also, the simulation has to be transitioned to include multiple layers. This would enable the study of the effects that the interaction between layers has on the properties of the sintered bed.1770ConclusionThis paper presents densification results which were derived from isothermal heating of copper nanoparticles. The data shows that after 10 minutes of heating at 450˚C the density value starts to reach the maximum steady state value. As heating temperature increases, the amount of time the nanoparticles take to reach steady state decreases. The densification curves obtained from the experimental data are calibrated to match sintering simulation results and it is seen that the simulation trend is in good agreement with the experimental data. The diffusion constants are mapped to physical units and are compared to bulk properties from literature.

The diffusion constants are mapped to physical units and are compared to bulk properties from literature. This comparison shows that the surface diffusion coefficient is higher for the copper nanoparticles than those reported for bulk copper.AcknowledgementsThis material is based upon work supported by the National Science Foundation under Grant No. 1728313. The authors would like to acknowledge Texas Advanced Computing Center (TACC) for the Supercomputing resources, and NXP Semiconductors for the financial support provided.References[1] Roy, N. K., Foong, C. S., and Cullinan, M. A., 2016, “Design of a Micro-scale Selective Laser Sintering System,” 2016 Annual International Solid Freeform Fabrication Symposium.[2] Dibua, O. G., Yuksel, A., Roy, N. K., Foong, C. S., and Cullinan, M., 2017, “Modelling Nanoparticle Sintering in a Microscale Selective Laser Sintering Process,” 2018 Annual International Solid Freeform Fabrication Syposium.[3] Khondoker, M. A. H., Mun, S.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEXPERIMENTAL CHARACTERIZATION OF A DIRECT METAL DEPOSITED COBALT-BASED ALLOY ON TOOL STEEL FOR COMPONENT REPAIRXinchang Zhang*, Tan Pan*, Wei Li*, Frank Liou*Department of Mechanical & Aerospace Engineering, Missouri University of Science and Technology, Rolla, Missouri, 65409, USAAbstractDie casting dies made of tool steel is subject to impact, abrasion and cyclic thermo- mechanical loading that delivers damage such as wear, corrosion, and cracking. To repair such defects, materials enveloping the damage need to be machined and refilled. In this study, V-shape defects with varied sidewall inclination angles were prepared on H13 tool steel substrates and refilled with cobalt-based alloy using direct metal deposition (DMD) for superior hardness and wear resistance.

The microstructure of rebuild samples was characterized using an optical microscope (OM) and scanning electron microscope (SEM). Elemental distribution from the substrate to deposits was analyzed using energy dispersive spectrometry (EDS). Mechanical properties of repaired samples were evaluated by tensile test and microhardness measurement. Fracture mechanism in tensile testing was analyzed by observing the fracture surface. The experiment reveals that V-shape defects with sidewall beyond certain angles can be successfully remanufactured. The deposits were fully dense and free of defects. The microstructure and tensile test confirm the solid bonding along the interface. The tensile test shows the mean ultimate tensile strength (UTS) of repaired samples is approximated 620 MPa, where samples fractured at the deposits region.

The tensile test shows the mean ultimate tensile strength (UTS) of repaired samples is approximated 620 MPa, where samples fractured at the deposits region. Hardness measurement reveals the hardness of deposits is around 810 HV which is much higher than that of the substrate.Keywords: Direct metal deposition; Repair; Co-based alloy; Tool steelIntroductionLaser-aided Direct Metal Deposition (DMD) is a relatively new technique but has shown great applications in the fields of part fabrication [1], [2], surface coating for improved wear and corrosion resistance [3], [4], joining different materials [5]–[7], deposit high-entropy alloys [8] and functionally gradient material [9]–[11] and part remanufacturing [12]–[17]. Depositing or cladding hard surfacing alloys on components through DMD process can modify the surface properties to greatly enhance hardness, wear and corrosion resistance, therefore, extending the service life of metallic components.

When defects were observed on worn components, materials with better failure resistance can be welded in the damaged area to assure that the remanufactured parts will last a long time.Cladding hard surfacing alloys such as Ni-based, Co-based alloys and WC composite has been widely investigated. Paul et. al. reported in [18] that they are able to produce fully dense and crack-free W-C-Co coatings on low carbon steel using pulsed Nd: YAG laser with an excellent bi- material bonding and much higher hardness. Zhong et. al. investigated the microstructure evolution during laser cladding of mixture of Stellite 6 and WC powder. In the coating of W-C-Co alloys on1451medium carbon steel [19], the researchers observed some defects like pores and cracks in the clad layers mainly owing to gas entrapped in clad layers and stress concentration. Spherotene tungsten carbides were deposited on low carbon steel in [20].

Spherotene tungsten carbides were deposited on low carbon steel in [20]. The analysis exhibits a uniform distribution of hard particles in coatings with a hardness around 1000-1500 HV. Investigators in [20] coated 1.0 – 3.3 mm thick Co-based coatings on cast iron by optimizing processing parameters. The result reveals crack-free and good mechanical performance deposits.Defects such as cracks, dents, pitting are commonly observed in worn components. In order to repair such defects, it is usually required to machine a slot or groove to remove materials around such damage to reveal a regular shape. In machining the slot, the sidewall inclination angle of the slot should be carefully determined. This is because the sidewall tilt angle may affect the bonding condition of filler material and substrate. A small tilt angle assures the accessibility of the target area but may result in much material removal.

A small tilt angle assures the accessibility of the target area but may result in much material removal. Because the removed material needs to be rebuilt through deposition, a large amount of material removal is not suggested. However, if the sidewall tilt angle is large, the bonding between deposits and substrate at the vertical surface cannot be guaranteed. Therefore, the slot sidewall inclination angle should be studied so that less material will be removed to get rid of defects and assure the bi-material bonding is solid.This paper has two objectives. The first is to investigate the feasibility of direct metal deposition of a Co-based alloy Wallex 50 on tool steel for repair purpose. Besides, this papers aims to study the influence of sidewall inclination angle of defects on the properties of repaired coupons. In order to perform repair process, V-shaped defects with sidewall inclination angle of 45°, 75° and 90° were prepared on H13 tool steel substrates.

In order to perform repair process, V-shaped defects with sidewall inclination angle of 45°, 75° and 90° were prepared on H13 tool steel substrates. The missing volume of each substrate is reconstructed by scanning the damaged part using a structured-light 3D scanner. After that, the missing geometry was sliced into layers to generate a raster deposition tool path. Wallex 50 material was added to the damaged region to regain the missing geometry. In order to test the repair quality, macrostructure and microstructure characterization, EDS and XRD analysis, tensile test and microhardness measurement were performed on the repaired samples.Experimental condition and procedureMaterial and preparationH13 tool steel is usually used for manufacturing die casting dies owing to its great thermal fatigue cracking resistance. Therefore, H13 tool steel plate with dimensions of 25 mm × 10 mm × 15 mm was selected as the substrate material.

Therefore, H13 tool steel plate with dimensions of 25 mm × 10 mm × 15 mm was selected as the substrate material. In order to perform repair experiments, several V- shaped grooves with different sidewall inclination angles were machined on the substrates. Three sidewall inclination angles of 45°, 75°, 90°, were chosen which are depicted in Fig. 1a-c. The depth of each groove is 7 mm and the width of defects at the bottom is 5 mm. The chemical compositions of H13 tool steel are listed in Table 1.Wallex 50 is chosen as the filler material because it is a Co-based hard surfacing alloy with excellent corrosion resistance and low coefficient-of-friction, providing good metal-to-metal wear protection. Wallex 50 has high contents of Co, Ni, Cr, and W to guarantee a high hardness as well as good corrosive resistance. Wallex 50 can be coated on easy-to-wear surfaces of H13 tool steel dies to enhance its resistance to wear.

Wallex 50 can be coated on easy-to-wear surfaces of H13 tool steel dies to enhance its resistance to wear. Therefore, this paper tries to evaluate the repair quality of1452using Wallex 50 as a deposition material. The chemical compositions of Wallex 50 are collected in Table 1. Wallex 50 alloys powder was characterized to analyze particle shape and size distribution. An SEM micrograph of Wallex 50 alloys powder was obtained and presented in Fig. 2a. The image reveals that most powder particles are spherical although a few of them have irregular shapes. Particle size distribution analyzed using ImageJ shows the average particle diameter is 65 µm (Fig. 2b).(a) (b) (c)Fig.

Particle size distribution analyzed using ImageJ shows the average particle diameter is 65 µm (Fig. 2b).(a) (b) (c)Fig. 1 Dimensions of the prepared substrates with (a) 45°, (b) 75° and (c) 90° sidewall inclination angleMaterials H13 tool steel Wallex 50C0.40.8Table 1 Chemical composition of the target materials (wt%) NiMnSiCrMoVW0.41.05.251.351.02.75191810B3.4FeCoBal.1.0 Bal.(a) (b)Fig. 2 SEM micrograph (a) and particle size distribution (b) of Wallex 50 powderExperimental setupWhen V-shaped defects were prepared on the substrate, it is necessary to obtain the geometry of the missing volume to generate additive tool path. This deposition tool path directly determines the shape of restored parts and has great effects on the mechanical properties of deposits. To make sure the acquired dimensions of the refill volume is accurate, an RE-based damage extraction process was conducted.

To make sure the acquired dimensions of the refill volume is accurate, an RE-based damage extraction process was conducted. Each substrate was scanned using a high accuracy1453structured-light metrology 3D scanner (OptimScan 5M from Shining 3D Tech. Co. Ltd., Fig. 3a) to reconstruct its three-dimensional model. The principle of the structured-light scanner is shown in Fig. 3a where a pattern with stripes was projected on an object by the structured-light projector and two CCD cameras on the left and right of the scanner capture the distortions of the pattern. The distortion of pattern reveals the three-dimensional geometry of the object.

The distortion of pattern reveals the three-dimensional geometry of the object. A complete scan of the part requires scanning of the geometry in different orientations and therefore, indexing targets were randomly marked on the part so that multiple scans can be registered into a single model.Substrates with different sidewall inclination angles were scanned using the configuration and the scanned point cloud was presented in Fig. 3b. Defining the point set in the damaged region is simply conducted by calculating the distance between each point to the top surface. The points on peripheral surfaces around the substrate are not considered as damage. After damaged point searching, the missing geometry can be defined, which is depicted in Fig. 3c. One can see in Fig. 3c that the damage for the sample with 90° sidewall defects was not completely scanned because the vertical sidewall combined with narrow slot blocks the two cameras from capturing projected pattern simultaneously.

However, this uncomplete scanned point data has no effect on the additive tool path generation because the tool path generation algorithm relies on the convex-hull of the data set, which is described below.After the damage was extracted, tool path was generated based on damaged point cloud utilizing the algorithm introduced in [21]. The damaged point set was sliced into 11 layers with a layer thickness of 0.6 mm. For each layer, deposition tool path was planned which comprises outline contour and raster infill pattern, indicating laser moves along the enveloping boundary and then along the zigzag pattern to acquire fully dense deposits. The tool path for three scenarios was shown in Fig. 4a-c.After additive tool path was acquired, the damaged substrates were loaded to DMD system for material deposition. The DMD cell employed in this study (Fig.

4a-c.After additive tool path was acquired, the damaged substrates were loaded to DMD system for material deposition. The DMD cell employed in this study (Fig. 5) includes a 1000 W peak power continuous wave fiber laser, gas-assisted blown powder feeder (Model 1200, Bay State Surface Technologies, Inc.), 3-axis work table and Argon gas feeding unit. The laser has a beam diameter of 1.8 mm. The laser beam is tilted while the powder feed nozzle is vertical with a stand- off distance of 10 mm. The welding was performed in argon gas environment to keep weld beads from oxidization. The processing parameters were summarized in Table 2.(a)1454(b)(c) Fig. 3 Model reconstruction for repair volume determination; (a) 3D scanning setup; (b) Reconstructed point cloud data; (c) Extracted damaged point set(a) (b) (c) Fig. 4 Repair tool path for substrates including defects with (a) 45°, (b) 75° and (c) 90° sidewall inclination angleFig.

4 Repair tool path for substrates including defects with (a) 45°, (b) 75° and (c) 90° sidewall inclination angleFig. 5 DMD system and setupTable 2 Component repair processing parameters Scan speed Layer thickness (mm/min) (mm) 220 0.6Laser power (W) Powder flow rate(g/min) 3.2350Overlap0.51455Sample preparation and material characterizationRepaired samples were sliced into layers using Hansvedt electric discharge machine (EDM) and mounted on Bakelite using Simplimet 1000 mounting equipment. Specimens were ground using abrasive papers from 120 Silicon Carbide Grid to 1200 Grid and finally polished using 0.05 microns silica suspension and then ultrasonically cleaned. The prepared samples were macroscopic analyzed using HIROX KH-8700 optical microscope to study the fusion boundary between refilled material and substrates. After that, the microstructure of deposits near the interface was revealed using Hitachi S4700 Scanning Electron Microscope (SEM).

After that, the microstructure of deposits near the interface was revealed using Hitachi S4700 Scanning Electron Microscope (SEM). EDS line scan analysis was performed from the deposits to substrates to reveal elemental distribution.In order to test the mechanical properties of the repair coupons, tensile test and hardness measurement were conducted. For tensile test, the samples were cut into thin layers with a thickness of 1 mm as shown in Fig. 6. Tensile specimens were cut from each thin slices using EDM. The dimensions of tensile specimens are shown in Fig. 6c. The position and orientation of tensile samples were depicted in Fig. 6a-b. Each tensile specimen consists of deposits and substrate, with interface locating at the middle of the specimen. The tensile test was conducted using Instron tester (Model 3300) with a crosshead speed of 0.015 mm/min (Fig. 6d). Ultimate tensile strength was obtained and then, tensile fracture surface was analyzed.

6d). Ultimate tensile strength was obtained and then, tensile fracture surface was analyzed. Vickers hardness was measured using Struers Duramin hardness tester with a press load of 9.81 N and dwell time of 10s.(a) (b)(c) (d) Fig. 6 Tensile test specimen preparation and setup; Tensile specimen over the sectioned repaired substrates with 45° (a) and 75° (b) sidewall inclination angle damage; (c) Dimensions of the tensile test specimen (Dimensions in mm, 1mm in thickness); (d) Tensile test setup1456Results and discussionVisual and MacrostructureThe overview of the repaired samples is shown in Fig. 7. It seems that missing volumes on all substrates were refilled successfully. However, by deeper visual observation, one can see the refilled region on the sample with 90° sidewall defects was not fully dense. A big hole was observed slightly at the right portion of deposits.

However, by deeper visual observation, one can see the refilled region on the sample with 90° sidewall defects was not fully dense. A big hole was observed slightly at the right portion of deposits. This is because when the laser was melting the materials on the side vertical surface, filler material was not able to bond with the substrate and fell down, which caused the hollow space inside the deposits. This is validated by observing the cross-sections of samples discussed below.The optical micrographs of cross-sections of the repaired samples are depicted in Fig. 8a- c, which has sidewall inclination angle of 45°, 75°, and 90°, respectively. Micrographs were taken at the boundary of deposits and substrate. One can realize that the bonding line is free of defects for 45° and 75° samples, while a large number of porosities and lack of fusion were observed in the sample with 90° tilt sidewall damage. Those defects were almost entirely located along the sidewall boundaries.

Those defects were almost entirely located along the sidewall boundaries. The bottom region of the 90° tilt sidewall sample, however, is well bonded. This phenomenon is expected because, for a vertical sidewall, the laser cannot continuously melt the materials at the vertical side of the defects but can still effectively melt materials on the bottom region. Thus, melt poor cannot be successfully formed at vertical regions and filler materials were not able to be melt and deposited. This situation might be addressed by using 5-axis DMD system by rotating the substrate so that the sidewall is relatively accessible to the laser beam. However, this is not always working if the defects are deep and narrow. Besides, toolpath planning for 5- axis DMD system is much complicated and only desired if really required.(a) (b) (c) Fig. 7 Overview of the repair samples with 45° (a), 75° (b) and 90° (c) sidewall inclination angle1457Fig.

7 Overview of the repair samples with 45° (a), 75° (b) and 90° (c) sidewall inclination angle1457Fig. 8 Optical micrographs of cross-sections of the repaired samples with 45° (a), 75° (b) and 90° (c) sidewall inclination angle damageMicrostructure characterization and EDS analysisBecause the defective substrate with 90° sidewall inclination angle was failing to repair, only samples with 45° and 75° sidewall damage were further analyzed in terms of microstructure and mechanical properties. Microstructure images were taken at the boundary of deposits and substrates, as shown in Fig. 9. The images distinctly show the bi-material boundary line. The microstructure changed sharply from the substrate to deposits. This is no smooth transitional region between two materials. By observation, one can see there are no defects such as cracks, pores, or delamination observed along the interface, confirming the solid bonding condition.

By observation, one can see there are no defects such as cracks, pores, or delamination observed along the interface, confirming the solid bonding condition. The SEM images reveal that deposits were metallurgically bonded with the substrate. The tensile test in the following of this paper also validated the sound bonding.The SEM micrographs in Fig. 9 also shows that the microstructure of deposits near the interface mostly consists of columnar structure stretched towards to the center of deposits. Those grains were growing parallel to the heat flow direction during solidification. When starting to deposit the first few layers, the damaged substrate was at room temperature. The laser melted a tiny amount of material on the substrate, forming a melt pool, and when materials were injected into such region and laser moved away, the cooling rate at such area is extremely high, which caused the grain to grow in the columnar direction.

The cooling is so fast that leaves not sufficient time for grains to form secondary dendrites. As materials were deposited layer by layer, the microstructure of deposits gradually changed to dendrite with interdendritic eutectics. This results from the dropped cooling rate when depositing the above layers.The microstructure of top layers of as-deposited Wallex 50 alloy is presented in Fig. 9c, which consists of mostly interdendritic eutectics. It was observed that the gray region was enveloped by bright phases. EDS spectrum shows the bright region in Fig. 9c was rich in Cr and W and the gray area is dominated by Co, Ni, and Si.EDS line scan was performed on the repaired samples to study elemental composition and distribution. The obtained data are plotted in Fig. 10. The scan was started from the deposits and ended in the region of the substrate, passing through the bi-material interface. Major elements of Co and Fe were quantified along this path.

10. The scan was started from the deposits and ended in the region of the substrate, passing through the bi-material interface. Major elements of Co and Fe were quantified along this path. Dwell time for each scanned point was 200 ms.1458EDS line scan exhibits the major elements like Co and Fe were gradually changed when passing the interface. This is because, when a melt poor was created on the substrate, Co-based alloy Wallex 50 was filled in the melt poor and mixed with the melted H13 tool steel material. Such mixed regions caused the elemental distribution changing near the interface. Relatively large amount of Fe exists in the first few layers of deposits, showing the dilution of Fe into Wallex 50. Therefore, Fe contributes to the phase transformation over the interface. The EDS line scan result confirms that excellent metallurgical fusion was created through the interface.(a) (b)(c) Fig.

Therefore, Fe contributes to the phase transformation over the interface. The EDS line scan result confirms that excellent metallurgical fusion was created through the interface.(a) (b)(c) Fig. 9 SEM micrographs of the boundary area of repaired samples with 45° (a) and 75° (b) sidewall damage; (c) Micrographs of top layers of the repaired sample with 45° sidewall damage1459(a) (b)Fig. 10 EDS line scan analysis results for samples with 45° (a) and 75° (b) damageMechanical propertiesTensile behavior and tensile fracture surfaceThe tensile stress-strain curves for two repaired samples were plotted in Fig. 11 and the Ultimate Tensile Strength (UTS) is summarized in Table 3. The tensile stress-strain curves reveal that the tensile stress increased with the increase of tensile strain to a maximum stress of about 620 MPa, and then, the tensile samples fractured suddenly and the stress decreased rapidly. The curves didn’t reveal yielding phenomenon during the tensile testing.

The curves didn’t reveal yielding phenomenon during the tensile testing. The ductility of tensile samples was extremely low. Further investigation reveals all samples fractured at the as-deposited Wallex 50 region. Therefore, the as-deposited Wallex 50 alloy shows brittle failure mechanism.The average UTS of the repaired samples with 45° and 75° V-shaped defects is approximately 618 MPa and 624 MPa, respectively. The tensile testing shows the influence of sidewall inclination angle on the UTS is barely recognizable. There is a relatively big variation of UTS in the tested result. This variation was possibly attributed to the unevenly located defects in the deposited regions. Some micropores in deposits may affect the UTS and result in such standard deviation.The micrographs of tensile fracture surface are presented in Fig. 12a-c. The overview image reveals a relatively flat surface which is perpendicular to the tensile loading direction. Magnified view in Fig.

12a-c. The overview image reveals a relatively flat surface which is perpendicular to the tensile loading direction. Magnified view in Fig. 12b-d shows microscopically unsmooth areas. Microcracks were observed at such magnified views. There is a very limited population of voids and dimples, indicating the fractured region is very brittle. This is because the Wallex 50 is a Co-based alloy and also consist of a relatively large number of hard elements including Cr, W, Si. Cr, W, and Si can form a very hard phase that results in the brittle failure mechanism of as-deposited Wallex 50 during tensile testing.1460(a) (b) Fig. 11 Tensile stress-strain curves for repaired samples with 45° U-shaped defects (a) and 75° U-shaped defects (b)Table 3 UTS obtained from tensile test45°75°Specimen 1 2 3 4 Average S.D. 1 2 3 4 Average S.D.UTS (MPa) 623.06 612.12 610.00 629.46 618.66 9.19 646.22 656.36 602.61 591.73 624.23 31.831461Fig.

1 2 3 4 Average S.D.UTS (MPa) 623.06 612.12 610.00 629.46 618.66 9.19 646.22 656.36 602.61 591.73 624.23 31.831461Fig. 12 Tensile fracture morphologies; (a) Overview of the fracture surface; (b) Magnified view of area 1 in (a); (c) Magnified view of area 2 in (a); (d) Magnified view of area 3 in (a)Microhardness measurementsVickers hardness was measured on the cross-sections of the repaired samples starting from the top of deposits to the substrate. The step between two adjacent indentations is 0.4 mm. The Vickers hardness tested on two samples was plotted in Fig. 13.A first observation is that the hardness of deposits is much higher than the hardness of substrate. The hardness of the H13 substrate is approximately 210 HV, while the hardness of deposits is 810 HV, almost four times of the hardness of substrate. However, H13 tool steel used for die casting dies is in quenched and tempered condition and usually shows a hardness range of 480 -520 HV.

However, H13 tool steel used for die casting dies is in quenched and tempered condition and usually shows a hardness range of 480 -520 HV. The reason that the hardness of the substrate utilized in this study is only 210 is owing to the condition of the received H13 tool steel material. The as-received H13 substrate is in the annealed condition and has not been hardened, causing the low hardness. However, one can see that even the hardened H13 tool steel which condition is exactly used for casting dies is also much lower than the hardness of as-deposited Wallex 50.The sharply increased hardness on the deposited region is mainly caused by two reasons. The major reason is that Wallex 50 is a Cobalt-based Co-Ni-Cr-W alloy. The Cobalt-rich matrix with very hard phases results in the very high hardness. In addition, the cooling rate during material deposition is rapid, leading to the formation of fine microstructure, which also plays a role in the1462hardening.

In addition, the cooling rate during material deposition is rapid, leading to the formation of fine microstructure, which also plays a role in the1462hardening. The hardness at bi-material interface jumped from 250 HV to 800 HV rapidly. The transitional hardness values at interface indicate the dilution in the heat-affected area. The transitional range is approximately 1.5 mm.Fig. 13 Vickers hardness tested on the cross-section of the repaired samplesConclusionIn the current study, Co-based alloy Wallex 50 was evaluated through depositing on H13 tool steel substrate. In order to test the possibilities of repairing slots with different sidewall inclination angles, V-shaped defects with 45°, 75°, and 90° sidewall tilt angles were prepared. The tool path for material deposition was obtained through 3D scanning of the defective substrate and extracting the damaged area. Once the missing volume was obtained, it was sliced into layers to generate a raster deposition tool path.

Once the missing volume was obtained, it was sliced into layers to generate a raster deposition tool path. Wallex 50 particles were filled up the missing area to regain the geometry. To validate the repair quality, a number of tests were performed on the repaired samples.The macroscopic analysis confirms the samples with 45° and 75° damage were repaired successfully. For the sample with 90° damage, since the laser was not able to melt the material at the vertical area, filler material was not bonding well with the substrate, causing lack of fusion and a large number of pores. The microstructure of deposits near interface shows mostly columnar structure mainly due to the high cooling rate for the first few layers. As materials built up layer by layer, microstructure was changed to the equiaxed structure. Microstructure and EDS analysis confirm the forming of metallurgical bonding along the interface.Tensile testing of repaired samples reveals a UTS of approximately 620 MPa.

Microstructure and EDS analysis confirm the forming of metallurgical bonding along the interface.Tensile testing of repaired samples reveals a UTS of approximately 620 MPa. The influence of sidewall inclination angle of damage on the UTS is not observed. The fracture surface shows all samples fractured at the as-deposit Wallex 50 region and shows brittle failure mechanism. The hardness analysis shows deposits is much harder than the substrate. A homogeneous hardness of 810 HV was measured on deposits while the hardness of substrate is around 210 HV.AcknowledgmentThis project was supported by National Science Foundation Grants CMMI-1547042 and CMMI-1625736, and the Intelligent Systems Center, Center for Aerospace Manufacturing1463Technologies, and Material Research Center at Missouri S&T. Their financial support is greatly appreciated.Reference[1]E. C. Santos, M. Shiomi, K. Osakada, and T. Laoui, “Rapid manufacturing of metal components by laser forming,” Int. J. Mach.

This technique relies on a piezoelectrically-driven ultrasonic print- head that generates continuous streams of droplets from 45 mm orifices while operating in the 0.5 to 3.0 MHz frequency range. Unique to this new printing technique are the high frequency of operation, use of fluid cavity resonances to assist ejection and acoustic wave focusing to generate the pressure gradient required to form and eject droplets. Specifically, we found that peaks in the ejection quality corresponded to predicted device resonances. Our results indicate that the micromachined ultrasonic print-head is able to print fluids up to 3000 mN-s/m2, far above the typical printable range.1 Recent developments in inkjet printing methods have not kept pace with the evolution of printed materials, from inks deposited onto porous surfaces in the form of two-dimensional (2D) text or images to complex fluids used for the free-form fabrication of three-dimensional (3D), multilayer devices.

Inkjet printing has a number of demonstrated advantages over conventional additive manufacturing (AM) technologies (e.g., stereolithography and laser sintering); unfortunately, the print-heads used in AM and other high resolution printing machines are restricted to materials with a limited range of physical properties. For example, viscosity is typically limited to below 40 mN-s/m2 and surface tension to above 28 mN/m [1, 2].Introduction1.1 Ink-jet printing has the promise of large-scale, economical manufacturing for a wide variety of polymer and, potentially, other materials. AM processes that are not printing-based are restricted to processing of materials at a single point. Conversely, printing-based methods are termed ‘scalable’ because the cost of adding more nozzles to an array is low. In addition, printing processes enable straightforward fabrication of multi-material parts and devices by dedicating different nozzles to different materials [1-3].

In addition, printing processes enable straightforward fabrication of multi-material parts and devices by dedicating different nozzles to different materials [1-3]. Although a number of droplet generation techniques are used in practice (see Le [4] for a thorough review), most AM systems utilize piezoelectric actuation. For example, MicroFab Technologies, Inc. (www.microfab.com, Plano, TX), FUJIFILM Dimatix, Inc. (www.dimatix.com, Santa Clara, CA), and Xaar, PLC (www.xaar.co.uk, Cambridge, UK) market print-heads that are capable of producing 30–100 µm diameter droplets from large nozzle arrays driven at piezoelectric pumping frequencies between 2 and 20 kHz.Inkjet printing methods, materials and applications1Many printing-based AM systems are available. Solidscape, Inc. (formerly Sanders Prototype, Inc., www.solid-scape.com, Merrimack, NH) was first to commercialize a printer that utilized independent nozzles for wax-based part and support structure materials.

(formerly Sanders Prototype, Inc., www.solid-scape.com, Merrimack, NH) was first to commercialize a printer that utilized independent nozzles for wax-based part and support structure materials. 3D Systems (www.3dsystems.com, Rock Hill, SC) has introduced a number of printing-based machines (e.g., the Actua 2100 in 1997, and the ThermoJet, InVision and ProJet lines in 2001, 2003 and 2008, respectively) with ever-increasing array sizes that now exceed several hundred nozzles. Objet Geometries, Inc. (www.object.com, Billerica, MA) has introduced machines (e.g., the Alaris30, Connex500 and Eden family of printers) with several-hundred to thousands of nozzle print- heads. The ProJet line and the Objet systems define each cross-section of a part in a photopolymer, which is then cured by ultraviolet light between each printing pass.

The ProJet line and the Objet systems define each cross-section of a part in a photopolymer, which is then cured by ultraviolet light between each printing pass. The Connex500 permits additional design complexity by printing two different materials, the relative concentrations of which can be adjusted to achieve varying properties. Although manufacturing parts using printing-based RP methods has proven to be fast and inexpensive, photopolymer materials and wax-based polymers do not possess the mechanical properties desired in most production applications; therefore, their usefulness is limited to prototype parts.A wide variety of fluids (e.g., biomaterial suspensions, ceramic pastes and non-Newtonian polymers) are relevant to 3D printing applications. Demonstrating the ability to print simultaneously cells, cell aggregates and biodegradable supports has driven recent developments in tissue engineering [5-11].

Demonstrating the ability to print simultaneously cells, cell aggregates and biodegradable supports has driven recent developments in tissue engineering [5-11]. Inkjet printing of conductive polymers and colloidal semiconductors is particularly useful for manufacturing flexible circuits [12], organic light emitting diodes (OLEDs) [13, 14], polymer photovoltaics [15] and transistors [3, 16]. A thorough introduction to conductive polymers in embedded electronic circuits and structural parts is provided by MacDiarmid [17]. Ainsley et al. [18] have printed ceramic alumina slurries with up to 40 percent solids (by volume) using wax and kerosene to reduce paste viscosity.

Ainsley et al. [18] have printed ceramic alumina slurries with up to 40 percent solids (by volume) using wax and kerosene to reduce paste viscosity. Polyurethanes represent an important class of materials for production applications; however, relatively high viscosities (up to 10 N-s/m2) hinder printing of urethane-based resins in conventional inkjet systems.1.2 Micromachining as an enabling technology for printer development Micromachining techniques are particularly well-suited to fabrication of arrays of uniform micron-sized orifices. As a consequence, the evolution of inkjet printing has been greatly influenced by developments in the area of micro-electromechanical systems (MEMS). The gradual reduction in nozzle orifice size and, as a result, droplet diameter from approximately 50 µm to less than 5 µm is the most obvious result of progress in microfabrication.

The gradual reduction in nozzle orifice size and, as a result, droplet diameter from approximately 50 µm to less than 5 µm is the most obvious result of progress in microfabrication. For example, orifice plates have been created by electroforming a metal over a pattern of nonconductive photoresist islands, which dictate the location and size of 2–6 µm orifices (see for example [19]). Atomization from thin (tens of microns) electroformed nozzle plates is typically achieved by vibrating the plate at high frequencies (~10 kHz). Device actuation has also been achieved by using a pressurized reservoir to push liquid jets from an array of laser-drilled 2.5 µm diameter holes in a polymer film [20]. The jets subsequently broke up into droplets with a mean diameter (~2.9 µm) slightly larger than that of the orifice. Femtoliter-sized (~5 µm diameter) droplets were produced by de Heij et al. [21] and Yuan et al.

Femtoliter-sized (~5 µm diameter) droplets were produced by de Heij et al. [21] and Yuan et al. [22, 23] using bending mode piezoelectric transducers operating at up to 200 kHz to actuate arrays of 5 µm diameter nozzles micromachined in 10–20 µm thick silicon membranes using deep reactive ion etching (DRIE). Perçin et al. [24-29] relied on a micromachined flextensional ultrasound transducer to excite the axisymmetric resonant modes in a clamped circular plate and periodically elevate the pressure within a fluid reservoir. If the pressure was high enough to overcome fluid inertia and surface2tension, a droplet was ejected. Droplet ejection was demonstrated for a number of orifices including some as small as 4 µm in a 100 µm diameter membrane driven at 3.45 MHz [24]. Although MEMS-based droplet generators exhibit a high degree of control over orifice uniformity and location, their fabrication processes are typically complex.

Although MEMS-based droplet generators exhibit a high degree of control over orifice uniformity and location, their fabrication processes are typically complex. Further, device robustness may be an issue in printing high-viscosity liquids.1.3 Additive Manufacturing via Microarray Deposition (AMMD) We propose a novel manufacturing technique termed Additive Manufacturing via Microarray Deposition (AMMD) that relies on a MEMS-enabled print-head to form and eject liquid droplets from arrays of micro-scale (3–50 µm) nozzles [25, 26]. Ultrasonic actuation, resonant operation and acoustic wave focusing combine to enable efficient deposition of materials that conventional print-heads are unable to print. The high frequency of operation (0.5–3 MHz) and parallel (array) format make possible fast deposition of large volumes of material. Further, multi-material structures can potentially be created by utilizing different reservoirs driven by one or more piezoelectric transducers.

Further, multi-material structures can potentially be created by utilizing different reservoirs driven by one or more piezoelectric transducers. The piezoelectric transducer operating frequency is matched to specific cavity resonant frequencies in order to efficiently transfer power from the piezoelectric transducer to acoustic waves within the sample reservoir. The pressure gradient required to print even high-viscosity materials is realized by exploiting the acoustic wave focusing properties of liquid horn structures. We demonstrate the ejection of liquids with a range of surface tensions (~25–73 mN/m) and viscosities (0.7–3000 mN-s/m2); a qualitative judgment termed ‘jettability’ is used to compare ejection of different liquids (e.g., water, glycerol-water mixtures and a urethane-based photopolymer resin).Figure 1. Schematic of the AMMD printer (not to scale).

Schematic of the AMMD printer (not to scale). a) Two-fluid multiplexed printer assembly including ejector microarray with two orifice sizes and b) cross-section of the print-head assembly.2 Print-Head Design The AMMD printer schematic shown in Fig. 1 illustrates its most basic multiplexed embodiment with two individually-actuated compartments. The device comprises a piezoelectric transducer for ultrasound wave generation, a reservoir for storing materials to be printed and a set of acoustic horn structures, terminated by microscopic orifices, which form a nozzle array for focused application of mechanical stimuli at the nozzle tips. Each compartment is loaded with the different fluids needed for fabrication of a multi-material part. When either chamber of the device is driven at a particular resonant frequency of the liquid horn structures, acoustic waves within the sample chamber are focused by the horns. The two sets of nozzles shown in Fig.

The two sets of nozzles shown in Fig. 1 are also terminated by two different orifice sizes allowing for easier implementation of multi-scale3part manufacturing. The realized ejector microarray consists of a rectangular fluid reservoir sandwiched between a lead zirconate titanate (PZT-8) piezoelectric ceramic (P880, American Piezo Ceramics, Inc.) and a silicon cover plate into which are micromachined. The pyramidal shape of the nozzles, which is suitable for focusing acoustic waves, is readily fabricated via anisotropic potassium hydroxide (KOH) wet etching of lithographically patterned [100]-oriented single crystal silicon.theejectornozzles3 Methodology3.1 Microarray fabrication A detailed description of ejector microarray is provided by fabrication Meacham et al. [26]. A simplified version of that process was used to create pyramidal nozzles in Si microarrays for use the experiments described herein.

[26]. A simplified version of that process was used to create pyramidal nozzles in Si microarrays for use the experiments described herein. In brief, the process consisted of four steps: (1) silicon nitride (Si3N4) was deposited over the front and back sides of a [100]-oriented single crystal Si wafer, (2) an array of squares was etched through the Si3N4 on one side of the wafer using a reactive ion etch (RIE) (SLR Series Dual ICP, Plasma-Therm), (3) a potassium hydroxide (KOH) solution (45 weight- percent at 75 °C) (VWR, Westchester, PA) was used to anisotropically etch pyramidal nozzles into the Si wafer, and (4) the remaining Si3N4 was removed using buffered oxide etchant (BOE) (VWR, Westchester, PA). The Si3N4 layer served as a protective mask during the anisotropic wet etch of the Si, and the square pattern controlled the size of the resulting pyramids.

The Si3N4 layer served as a protective mask during the anisotropic wet etch of the Si, and the square pattern controlled the size of the resulting pyramids. The fabrication process parameters were adjusted such that each nozzle terminated in a microscopic square orifice.in silicon(Si)inFigure 2. Scanning electron microscope (SEM) images of silicon (Si) microarrays used in the ejection experiments, with 45 µm square orifices.Figure 2 provides scanning electron microscope (SEM) images of a realized ejector microarray. The shape of the nozzles was dictated directly by the KOH wet etch, which yielded 45 µm side length square orifices on the opposite side of the wafer. The cross-section of the nozzle clearly illustrates the geometry of the slanted sidewalls, which lie at an angle of 54.74° to the (100) surface of the Si wafer.

The cross-section of the nozzle clearly illustrates the geometry of the slanted sidewalls, which lie at an angle of 54.74° to the (100) surface of the Si wafer. The silicon plate measures 24 mm on a side with 400 nozzles located in a 20×20 array on an approximately 0.75 mm pitch.3.2 Experimental setup Evaluation of the AMMD technique was performed using an experimental setup similar to that previously reported for high-resolution visualization of the ejection process [25, 26]. A schematic of the fully-assembled test fixture is shown in Fig. 3a-b. The silicon ejector microarray was bonded to a plastic support plate with epoxy. The plastic plate was created in Renshape SL- (Huntsman Advanced Materials Americas, Los Angeles, CA) using a 7510resin4stereolithography process (Viper Si2 SLA Machine, 3D Systems, Valencia, CA).

The 2 mm thick piezoelectric transducer (P880, APC International, Mackeyville, PA) was likewise affixed to an aluminum support block, which allowed for efficient heat removal during high-power-input operation. A circular Teflon® gasket was placed between the plastic plate and aluminum block to define the sample reservoir between the microarray and piezoelectric transducer. For all experiments, the height of the reservoir was between 1 and 2 mm. A large barrel syringe was used to fill the reservoir prior to operation. The piezoelectric transducer was driven by a sinusoidal AC voltage signal that was supplied by a function generator (Model DS345, Stanford Research Systems, Sunnyvale, CA). The signal was amplified using an RF amplifier (Model AG1020, T&C Power Conversion, Rochester, NY). The function generator provided control of the frequency f and amplitude Vfg of the driving signal, as well as the number of pulses n and pulse frequency (period) fd.c.

The function generator provided control of the frequency f and amplitude Vfg of the driving signal, as well as the number of pulses n and pulse frequency (period) fd.c. (Pd.c.) when operating at less than 100% duty cycle. Signal gain amp. was set using the amplifier. A type T thermocouple (Omega, Stamford, CT) was used to continuously monitor the temperature on the back surface of the transducer.Successful ejection of high-viscosity liquids (e.g., glycerol-water mixtures and photopolymer resins) required significant power input to the piezoelectric transducer. Under these conditions, high electrical current and physical displacement of the piezoelectric transducer led to increased resistive and frictional losses.

Under these conditions, high electrical current and physical displacement of the piezoelectric transducer led to increased resistive and frictional losses. During ejection of photopolymer resins, the entire test fixture was lowered into a temperature-controlled water bath to avoid excessive heating and maintain the desired operating temperature within the sample reservoir.3.3 Sample preparation and property determination Although many fluids have been ejected using our printing system, we will focus on aqueous glycerol mixtures and a representative photopolymer resin with urethane oligomers (Albany International Corporation, Albany, NY) for this paper. Aqueous glycerol mixtures with a range of fluid properties (e.g., surface tension and viscosity), as well as a photopolymer resin that is representative of materials of interest in AM, were used as operating fluids.

Because glycerol is completely miscible with water, it is well-suited to investigation of viscous effects in fluid flow problems. Glycerol-water mixtures of 0, 65, 85, 90, 92, 96 and 100 weight-percent were prepared from water and 99.7% pure glycerol (Fisher Scientific, Fair Lawn, NJ). Although surface tension σ plays an important role in the ejection process, its variation (~25–73 mN/m) was small relative to that of the dynamic viscosity µ (~0.7–3000 mN-s/m2) for these mixtures over the range of experimental conditions (i.e., temperature T < 60 °C).

Finally, the fluid compositions that were required to achieve a good spread in viscosity and surface tension data also exhibited a range of sound transmission speed c (~1400–1950 m/s), which plays a significant role in the optimal operating frequency of a print-head loaded with a given fluid.Glycerol-water mixtures have been thoroughly characterized in the literature; therefore, the properties of the above mixture compositions were taken from tabulated data (surface tension and speed of sound) and an empirical formula derived from multiple property databases (viscosity). Viscosities at temperature varied from 15 to 625 mPa-s. At printing temperature of 40 °C the viscosity of pure glycerin was 326 mPa-s.5Figure 3. Experimental setup: a) schematic of the test fixture highlighting integral components and b) detail providing nominal dimensions of the microarray assembly. Jettability: the still photographs define a qualitative assessment tool for evaluating ejection capability.

Jettability: the still photographs define a qualitative assessment tool for evaluating ejection capability. Scale bars are 10 cm.Because no information is available in the literature, the fluid properties (e.g., surface tension, speed of sound and viscosity) of the urethane photopolymer were measured experimentally, as functions of temperature when possible. A modified version of the drop-weight-volume method originally introduced by Harkins and Brown [36, 37] was used to estimate the surface tension. Estimated surface tension values produced by this method included water (ρ = 1000 kg/m3) σ = 70.1 mN/m, glycerol (ρ = 1260 kg/m3) σ = 58.7 mN/m, and polydimethylsiloxane (PDMS) (ρ6= 970 kg/m3) was σ = 21.6 mN/m. Considering the error associated with the determination of the average droplet volume, these values are extremely accurate. The surface tension of the photopolymer resin was found to be 29 mN/m.

Considering the error associated with the determination of the average droplet volume, these values are extremely accurate. The surface tension of the photopolymer resin was found to be 29 mN/m. Although, measurements were only conducted at room temperature, the results provide an estimate of the photopolymer resin surface tension, which lies within a range bounded by the other reference fluids.The speed of sound within the photopolymer resin was also determined using an in-house experimental setup. A plastic container that measured 11 mm on a side was filled with the sample liquid. The plastic container was then placed in a temperature controlled water bath. An ultrasonic transducer was driven at 7.5 MHz by a pulser-receiver (Panametrics 5072PR, Olympus NDT, Waltham, MA), which generated a sound wave that travelled through the liquid perpendicular to the container. The echo times for reflections from the front and back faces of the container were recorded by the receiver.

The echo times for reflections from the front and back faces of the container were recorded by the receiver. The distance traveled (22 mm) was then divided by the time difference to obtain the speed of sound within the photopolymer resin. The speed of sound as a function of temperature T varied almost linearly from 1525 at 25 °C to 1390 at 65 °C.A cone-plate rheometer (Physica MCR300, Anton-Paar GmbH, Graz, Germany) was used to evaluate the viscosity of the photopolymer resin. In order to confirm that the high-viscosity liquid was a Newtonian the viscosity was monitored while varying the shear rate γ at a constant temperature of 50 °C. The shear rate was increased from 0.01 to 100 s-1 in 20 steps and brought back to 0.01 s-1 in 10 steps. As indicated in Fig. 4a, the fluid retained its Newtonian character throughout the range of shear rates that were investigated.

As indicated in Fig. 4a, the fluid retained its Newtonian character throughout the range of shear rates that were investigated. Note that the deviation from Newtonian behavior observed at low shear rates was due to limitations of the rheometer and not to the itself. The viscosity was also fluid measured as a function of increasing and decreasing temperature between 25 and 65 °C. Figure 4b is a plot of the average viscosity of four data sets (two increasing and two decreasing) over the temperature range of interest. Note that the maximum viscosity error measurements was 1.5%. A 4th-order polynomial that was fit to the data is also shown in Fig. 4b. The reported viscosity for all subsequent experimental conditions was determined using this relationship.fluid,forallaverage3.4 Assessment of printingcapabilityWe evaluated the ability of the AMMD technique to print a particular fluid byFigure 4.

Viscosity µ for the urethane photopolymer resin: a) Viscosity as a function of shear rate γ and b) temperature T.7making a qualitative assessment of fluid atomization and comparing the observed performance to a reference scale. The still photographs shown in Fig. 3 were taken during water ejection under a range of operating conditions, which are provided below each photograph. The measured height of the plume and an estimate of the number of active nozzles were used to evaluate ejection quality on a scale of 1 to 10. The camera was unable to capture operation at the low end of the scale with steady, but weak, ejection from only 3 to 5 orifices; however, at peak performance, the device demonstrated jet ejection from more than 200 nozzles generating a large plume approximately 50 cm in height (see Fig.

3).4 Results and Discussion4.1 Jettability as a function of frequency The acoustic responses of water and fuel-loaded ejector microarrays have previously been investigated using the harmonic analysis package of the commercial finite element analysis code ANSYS [38, 39]. These simulations indicated which frequencies corresponded to optimal device operation from both an efficiency standpoint and to achieve maximum power transfer from the piezoelectric transducer to generated droplets; however, the increase in operating temperature associated with operation at high electrical current and physical displacement of the piezoelectric transducer near resonance was not addressed.

This phenomenon is particularly important for printing of high-viscosity fluids, which exhibit large changes in viscosity with only a slight increase in temperature.In order to determine the relationship between operating frequency and ejection quality for high- viscosity fluids, the ‘jettability’ Jab of glycerol-water mixtures with increasing weight-percent glycerol was evaluated. Figure 5 provides plots of Jab as a function of frequency for 65, 85, 90, 92, 96 and 100 weight-percent glycerol. For all cases, the device was operated at less than 100% duty cycle with a pulse frequency fd.c. of 500 Hz and a pulse count n of 200. The signal gain amp. was set to 100%, and the amplitude Vfg of the driving signal was adjusted in order to achieve an observable increase in ‘jettability’ at the optimal ejection frequency.

The signal gain amp. was set to 100%, and the amplitude Vfg of the driving signal was adjusted in order to achieve an observable increase in ‘jettability’ at the optimal ejection frequency. For example, Vfg = 1 V was adequate to achieve Jab = 9 for 65% glycerol at f = 950 kHz; whereas, Vfg = 10 V was necessary to achieve only Jab = 7.5 for 100% glycerol at the same operating frequency.In general, the results shown in Fig. 5 follow the rough guideline for device design previously reported by Meacham et al. [39]. Fluid cavity resonances fc are approximately related to the reservoir centerline height hc and the speed of sound c within the ejection fluid as fc ≈ m c/(2κ hc) where m is the particular resonant mode. The reservoir height must be adjusted by the factor κ to account for the actual path length between the face of the piezoelectric transducer and the pressure wave focal point.

The reservoir height must be adjusted by the factor κ to account for the actual path length between the face of the piezoelectric transducer and the pressure wave focal point. For the first and second cavity resonant modes, the correction factors were reported as 0.77 and 0.93, respectively [39]. For 65% glycerol (c = 1800 m/s), peaks in ‘jettability’ are found at frequencies of 550 and 950 kHz, which are very near the predicted resonant frequencies of 580 and 970 kHz (note height h is 2 mm). As the weight-percent glycerol increases, only a slight shift of the resonant frequencies to to higher frequencies is seen. Although the speed of sound in pure glycerol is 10% greater than that in 65 weight-percent glycerol in water at room temperature (approximately 1950 vs. 1800 m/s), it is possible that the observed increase in operating temperature with glycerol concentration mitigates this effect somewhat.8Figure 5.

1800 m/s), it is possible that the observed increase in operating temperature with glycerol concentration mitigates this effect somewhat.8Figure 5. Jettability Jab as a function of operating frequency f for glycerol-water mixtures with increasing weight-percent glycerol. For all cases, the device was operated at less than 100% duty cycle with a pulse frequency fd.c. of 500 Hz and a pulse count n of 200.9It is clear from Fig. 5 that increasing the viscosity of the operating fluid is detrimental to the quality of ejection. As the glycerol concentration is increased from 65 to 100%, with a corresponding increase in viscosity from 10 to 326 mN-s/m2, ‘jettability’ at the first resonant frequency decreases quickly until all activity ceases at above 96 weight-percent glycerol. Although ‘jettability’ does not vanish entirely at the second resonance, the frequency envelope over which ejection can be achieved is significantly reduced for higher-viscosity mixtures.

Although ‘jettability’ does not vanish entirely at the second resonance, the frequency envelope over which ejection can be achieved is significantly reduced for higher-viscosity mixtures. It is possible that the decrease in surface tension from 67 to 61.5 mN/m corresponding to the change in conditions represented by the data in Fig. 5 contributes to the drop in device performance; however, it is not likely that this has as significant an impact as the greater than one order-of- magnitude increase in viscosity.4.2 Jettability as a function of duty cycle Jettability was also investigated while varying the pulse count n of the signal driving the piezoelectric at a constant operating frequency f of 950 kHz and pulse frequency fd.c. of 500 Hz. Figure 6 is a summary of results for mixtures with 65, 85, 90, 92, 96 and 100 weight-percent glycerol. As before, the signal gain amp.

of 500 Hz. Figure 6 is a summary of results for mixtures with 65, 85, 90, 92, 96 and 100 weight-percent glycerol. As before, the signal gain amp. was set to 100%, and the amplitude Vfg of the driving signal was adjusted in order to achieve observable changes in ‘jettability’ as the pulse count was decreased from 100 (duty cycle = ~0.05) to zero. As expected, the general trend displayed in Fig. 6 is a gradual increase in ‘jettability’, which appears to approach a maximum value beyond n = 100. More interesting is the ejection behavior as the pulse count approaches zero; even at duty cycles as low as 0.0025 (n = 5), faint ejection was observed.

More interesting is the ejection behavior as the pulse count approaches zero; even at duty cycles as low as 0.0025 (n = 5), faint ejection was observed. This result indicates that a small number of cycles is required to establish a favorable pressure field fluid reservoir.for ejection withintheThe present system represents a driven, damped harmonic oscillator for which the quality factor Q of a resonance is expressed asQ =τωo 2(1)where τ is the rise time for pressure wave amplitude to reach 1 – 1/e = 0.63 of its steady state value and ωo is the angular frequency at resonance [40]. For high-Q systems, the quality factor can also be defined in terms of the full-width at half-maximum of the amplitudeQ =ωo ∆ω(2)where ∆ω is the angular frequency difference at half of the maximum signalFigure 6. Jettability Jab as a function of pulse count n for glycerol-water mixtures with increasing weight-percent glycerol.

Jettability Jab as a function of pulse count n for glycerol-water mixtures with increasing weight-percent glycerol. For all cases, the operating frequency f was 950 kHz, and the pulse frequency fd.c. was 500 Hz.10amplitude at resonance. Solving for τ in terms of frequencyτ=1 π∆f .(3)Based on Fig. 5 and results reported earlier by Meacham et al. [39], the frequency difference at half-maximum for a resonant frequency of 950 kHz is between 50 and 100 kHz. Substituting into equation (3), the rise time should be approximately 5 µs, which corresponds to five pulses of the signal driving the piezoelectric. This simple analysis provides an upper limit on the minimum number of pulses required to initiate ejection and indicates that the conclusions drawn from Fig. 6 are realistic.4.3 Maximum and sustained jettability In order to obtain the results shown in Figs.

6 are realistic.4.3 Maximum and sustained jettability In order to obtain the results shown in Figs. 5 and 6, the device was only operated for 5–10 s to evaluate the printing potential under the reported experimental conditions. Incorporating the microarray into a commercial print-head will require both demonstration of a high deposition rate, which is related to maximum achievable ‘jettability’, and continuous operation. As before, glycerol-water mixtures with 65, 85, 90, 92, 96 and 100 weight-percent glycerol were evaluated. Sustained ejection was only observed at the first resonance for 65% glycerol. At an operating frequency of 550 kHz (amp. = 100%, Vfg = 1 V, n = 100 and fd.c. = 500 Hz), ejection at Jab = 6 was achieved for 90 s at an operating temperature T of 35 C.

At an operating frequency of 550 kHz (amp. = 100%, Vfg = 1 V, n = 100 and fd.c. = 500 Hz), ejection at Jab = 6 was achieved for 90 s at an operating temperature T of 35 C. For all other mixtures at the first resonant frequency (~550 kHz), either sustained ejection was not possible, or the rise in operating temperature associated with prolonged operation was unacceptable.Devices driven at the second resonant frequency (~950 kHz) exhibited more stable operation. The ‘jettability’ of 65% glycerol remained steady at Jab = 8 for 86 s at T = 35 °C (amp. = 100%, Vfg = 1 V, n = 100 and fd.c. = 500 Hz). As the weight-percent glycerol was gradually increased from 85 to 90, 92, 96 and 100%, it was necessary to increase the amplitude Vfg of the signal driving the piezoelectric transducer from 1 to 5 V to achieve a reasonable ‘jettability’. Over this range of mixture compositions, the ‘jettability’ slowly dropped from 8 to 4.5.

Over this range of mixture compositions, the ‘jettability’ slowly dropped from 8 to 4.5. For all but one of these cases, ejection was sustained for more than 70 s. The decrease in performance is attributed in part to a corresponding reduction in pulse count, which was necessary to control the temperature; however, even at n = 40, the steady state operating temperature rose to as high as 50 °C for the mixtures under investigation. Based on these results (mixture compositions and operating temperatures), the maximum viscosity for sustained ejection was approximately 150 mN-s/m2. Finally, note that a ‘jettability’ of 10 (or the scale maximum from Fig. 3) was achieved for 65, 85, 90, 92, 96 and 100 weight-percent glycerol at operating temperatures of 35, 40, 56, 54, 63 and 70 °C, respectively, and a frequency of 950 kHz.4.4 Ejection of photopolymer resin The potential of AMMD for 3D inkjet printing was evaluated using the urethane photopolymer resin.

Based upon the measured speed of sound (between 1400 and 1450 m/s over the range of operating temperatures, 40−60 °C) and using the same geometry as before, the first two resonant frequencies were estimated to be 500 and 900 kHz. Stable ejection was not possible at either frequency; however, sustained ‘activity’ was observed over 50 kHz windows at approximately 510 and 920 kHz (see Fig. 7c). In order to achieve observable jet ejection, it was necessary to clean the front face of the ejector microarray with methanol, which was allowed to evaporate before switching on the amplifier. Ejection was short-lived (~0.5 s) as fluid would quickly begin to weep from the nozzle orifices and form a thin layer, which would continue to oscillate. Figure117 provides images of the experimental setup after cleaning, immediately after turning on the device, and after a period of weeping (10−15 s).

Figure117 provides images of the experimental setup after cleaning, immediately after turning on the device, and after a period of weeping (10−15 s). The inability to sustain ejection is not thought to result from the increase in fluid viscosity, but from the decrease in surface tension reported in the Methodology section above. Similar behavior has been reported for atomization of hydrocarbon fuels (e.g., kerosene and methanol) [39]. Deposition of pink fluid onto a white cloth was used to confirm that the short burst of ejection was indeed photopolymer resin (and not methanol). Although not particularly remarkable, these results are encouraging, and represent the first successful attempt to print a photopolymer resin with such a high viscosity (~3000 mN-s/m2).4.5 Printing Indicator (p.i.) Several investigators have identified a the dimensionless parameter ‘printing indicator’ (p.i.) that can be used to predict the printing potential of a given fluid [41]. The p.i.

The p.i. is defined as the ratio of the Reynolds number to the square root of the Weber number [p.i. = Re/We1/2 = (ρ r σ)1/2/µ] and is really just the better known the Ohnesorge number Oh. For successful printing, it is recommended that the p.i. be kept within a range of 1 to 10 [18]; however, as shown in Fig. 8, the proof- of-concept experiments AMMD conducted with glycerol-water mixtures and representative photopolymer resin fall outside of this range. Aside from a small number of data points taken while operating at relatively high temperature (> 60 °C) or at low glycerol concentration, the demonstrated ejection of glycerol-water mixtures occurred for fluid viscosities of greater than 80 mN- s/m2.

The fluid properties of the urethane photopolymer resin (low surface tension and high viscosity) suggest that it is to print, yet short-lived impossible droplet generation was achieved at surface tensions below 30 mN/m and viscosities up to 3 N-s/m2 (note the lower bound on the surface tension envelope in Fig. 8 is estimated based on the behavior of other fluids; surface tension of the photopolymer resin was not measured as a function of temperature). These preliminary results indicate that the ejector microarray is capable of printing fluids with much higher viscosities than conventional print-heads.termedinverse oftheFigure 7. Operation using urethane photopolymer resin: a) Picture of the ejector microarray after cleaning the front face with methanol, b) the microarray immediately after turning on the device (a short burst of ejection was observed lasting approximately 0.5 s), and c) the microarray after a period (~10-15 s) of weeping.

Note that in (c) the liquid is very active.12Figure 8. Printing indicator for successfully ejected fluids. Open circles indicate glycerol-water mixtures for which the operating temperature was above 60 °C. All other operating conditions were below 45 °C. Glycerol-water mixtures fall in a surface tension envelope between ~58 and 67 mN/m, while the surface tension of the photopolymer resin is only known to lie below 30 mN/m.5 Conclusions We have demonstrated a MEMS-enabled printing technique, which relies on ultrasonic actuation, resonant operation and acoustic wave focusing to produce micro-scale (~50 µm) droplets from a parallel array of pyramidal nozzles. Bulk piezoelectric actuation was used to drive ejection at between 0.5 and 1.0 MHz in order to characterize the acoustic response (and thus resonant behavior) of a 20×20 array of 45 µm side length square orifices etched in a 500 µm thick silicon plate and placed above a 1.5 mm high fluid reservoir.

Peaks in the ejection quality were found close to predicted values at frequencies of 550 and 950 kHz. Glycerol-water mixtures and a urethane-based photopolymer resin with a range of properties (surface tensions of ~25– 73 mN/m and viscosities of 0.7–3000 mN-s/m2) were successfully ejected, and the printing capability was evaluated using a qualitative judgment termed ‘jettability’.13The combination of operating parameters realized by the ejector microarray makes possible the deposition of high-viscosity materials that currently cannot be used in printing-based additive manufacturing processes. Because the fundamental physical limits on ejection of high-viscosity fluids are not well understood, the performance envelope of printing such fluids is only loosely defined; thus, development of appropriate ejection technologies has been impeded. Our future efforts will focus on improving device consistency and achieving sustained operation with high- viscosity (> 1000 mN-s/m2) materials.

Our future efforts will focus on improving device consistency and achieving sustained operation with high- viscosity (> 1000 mN-s/m2) materials. We will also characterize device operation for fluids with properties that lie between the extremes (low surface tension, high viscosity and high surface tension, low viscosity) presented here.6 Acknowledgements The authors would like to express our appreciation for financial support and use of the photopolymer resin with urethane oligomers courtesy of Albany International Research Corporation (Mansfield, MA). We would also like to thank Dr. Victor Breedveld (GT School of Chemical and Biomolecular Engineering, Complex Fluids Group) and Sarp Satir (GT, ME) for assistance in measuring the viscosity and speed of sound, respectively, of the photopolymer resin. All authors were employees at Georgia Tech during the research that led to this paper.References1.2.3.4.5.6.7.8.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEXPERIMENTAL CHARACTERIZATION OF THE MECHANICAL PROPERTIES OF 3D PRINTED BÉZIER-BASED LATTICE BEAMSA. Álvarez-Trejo*, E. Cuan-Urquizo*†, and A. Roman-Flores*†Tecnologico de Monterrey, School of Science and Engineering, Av. Eugenio Garza Sada 2501 Sur, Monterrey 64849, Mexico † Tecnologico de Monterrey, Institute of Advanced Materials for Sustainable Manufacturing, Av. Eugenio Garza Sada 2501 Sur, Monterrey 64849, Mexico AbstractArchitected materials are widely used in additive manufacturing to reduce weight. The controlled arrangement of material allows to tailor their mechanical properties by tuning their geometrical parameters. A parametrization based on cubic Bézier curves is employed here to generate lattice beams by changing the position of a free control point.

A parametrization based on cubic Bézier curves is employed here to generate lattice beams by changing the position of a free control point. Two topologies with the same volume fraction and base curve for the lattice constituent elements at different positions are studied and compared. Lattice beams are manufactured via Fused Filament Fabrication of polylactic acid. The effective stiffness and yield stress of these lattice beams is analyzed experimentally using three-point bending tests. Adjusting the control point location leads to tailoring the effective mechanical properties of the lattice beams. This methodology leads to the synthesis of architected topologies with customized mechanical properties.IntroductionCellular materials, also known as lattices or metamaterials, are controlled arrangements of material filling a defined volume. Some advantages of using cellular materials are minimal use of material, energy absorption, and high flexibility [1].

Some advantages of using cellular materials are minimal use of material, energy absorption, and high flexibility [1]. Cellular materials are used to manufacture lightweight parts, creating sandwich structures and applications for the transport sector [2, 3]. Given that the lattice-conforming elements are significantly smaller than the macro dimensions of the filled volume, cellular materials can be considered as homogeneous and described in terms of effective properties, which can be obtained experimentally as the properties of a solid material would [4]. These effective properties are a function of the lattice topology and volume fraction [5], i.e., the percentage of volume occupied by solid material. Then, the effective mechanical properties of a lattice, such as its stiffness, could be tailored by modifying parameters defining its topology.

Then, the effective mechanical properties of a lattice, such as its stiffness, could be tailored by modifying parameters defining its topology. Complex parametric curves such as NURBS [6], B-splines [7, 8], or Bézier curves [9] have been used to define the topology of lattices, where changing the position of the curve-defining control points modifies the effective mechanical properties of these lattices. An increased curvature in the lattice topology has shown to increase the flexibility of the structures [10].A controlled arrangement of material can be used to build beam-type structures, and their effective properties compared to those of equivalent homogeneous solid structures. The mechanical properties of these lattice beams have been studied in static loading conditions, such as in the study by Meng et al. [11], who analyzed the bending of sandwich panels with a three- dimensional kagome lattice core. Similarly, Wang et al.

[11], who analyzed the bending of sandwich panels with a three- dimensional kagome lattice core. Similarly, Wang et al. [12] created lattice beams and obtained their flexural rigidity in two bending planes. This mechanical characterization can also be done computationally, such as in the work by Álvarez-Trejo et al. [13], who studied the influence of1885shear on the in-plane deflection of hexagonal and re-entrant lattice beams, via numerical models. Lattice beams conformed of structured materials have various applications, such as the one developed by Eremeyev et al. [14], who proposed analytical and numerical models for the enriched buckling of a pantographic beam arrangement. Additionally, Gu et al.

[14], who proposed analytical and numerical models for the enriched buckling of a pantographic beam arrangement. Additionally, Gu et al. [15] used a porous core on a sandwich panel to induce a twist and increase the aerodynamic efficiency of a rotor blade.Few works focus on the experimental characterization of the mechanical properties of lattice beams; moreover, cellular topologies based on parametric curves have not been used to conform lattice beams. In this work, we focus on the out-of-plane flexural properties of Bézier- based lattice arrangements.MethodologyDesign and manufacturing of Bézier-based lattice beam samplesHere, two sets of Bézier-based lattices were generated from a base cubic Bézier curved segment. The definition of this curve is based on the methodology proposed in [9], but now controlling the position of the free control point by means of two independent variables: the distance r from the origin and the angle β describing its position.

See Figure 1a, where the cubic curve has fixed starting and ending control points (P0 and P3). The distance between these points is λ, and they are placed in the horizontal axis so that the midpoint of the cubic Bézier curve is coincident with x = λ/2. The position of a free control point P1 is then defined by the distance r and the angle β, as shown in Figure 1a. Control point P2 acts as a double mirror of P1 so that the cubic Bézier curve is symmetric. All Bézier curved segments were thickened to a total thickness t and arranged into two different topologies by rotating, flipping, and repeating the base curve. The two topologies studied here are labeled as single curve and double curve, shown in Figure 1b and Figure 1c, respectively.Figure 1. Lattice beam generation: (a) Base Bézier segment and defining geometric parameters. Two topologies are generated: (b) single curve and (c) double curve.

Lattice beam generation: (a) Base Bézier segment and defining geometric parameters. Two topologies are generated: (b) single curve and (c) double curve. An example of a 3D-printed sample is shown in (d).1886For the experimental design, a set of samples of the double curve topology (see Figure 1c) was manufactured, using four values of β: 30°, 45°, 60°, and 90°. For this first set, the distance r between the origin and control point P1 took the values r = 2 mm, r = 4 mm, and r = 6 mm. Finally, the unit cell length was set constant at λ = 10 mm. Samples of each topology are depicted in Table 1; r = 0 produces the square lattice, which is considered for comparison purposes for all values of β.A second set of samples was generated for the single curve topology (Figure 1b). In this case, the value of the angle β was fixed at β = 90° and the unit cell length was now set to λ = 5 mm, in order to match the volume fraction of the double curve topology.

In this case, the value of the angle β was fixed at β = 90° and the unit cell length was now set to λ = 5 mm, in order to match the volume fraction of the double curve topology. The distance r was set to r = 2 mm, so that r/λ = 0.4 as in one of the cases of the double curve topology.In both cases, the thickness of the curved elements was kept constant at t = 0.5 mm. This thickness was proposed based on the printer nozzle diameter, so that each cubic Bézier curve was printed in a single step. All CAD files were generated with SOLIDWORKS® 2021, using global variables and configurations for parameter variation inside a single CAD file. The generated geometries were then converted to independent STL files for 3D printing.Table 1.

The generated geometries were then converted to independent STL files for 3D printing.Table 1. Representative views of each fabricated topology.r/λ00.2Double curve 0.40.6Single curve 0.430°45°β60°90°Considering all the combinations in Table 1, lattice beams were generated with external dimensions, as length l = 120 mm, width w = 20 mm, and depth d = 4 mm. Samples were fabricated via Fused Filament Fabrication (FFF), using the Ultimaker® 2+ printer and white Ultimaker®1887polylactic acid (PLA). Ultimaker® Cura v4.10 was used to prepare the samples for printing. STL files were sliced with a layer height of 0.15 mm and considering a nozzle diameter of 0.4 mm. Samples were placed flat in the building plate, so that the direction of the depth d was the same as the building direction, which allowed to print without the need for supports or build plate adhesion.

Samples were placed at the build plate in the same orientation, with nine lattice beams per printing job. The printing speed was kept constant at 50 mm/s and print cooling was enabled. Three replicates of each design were printed.Mechanical testing of lattice beam samplesAccording to Euler-Bernoulli beam theory, the maximum deflection δ of a simply supported beam loaded at the midpoint is δ = FL3/(48EI), where F is the applied load, L the effective length of the beam, E the Young’s modulus, and I the second moment of area [16]. Given that the simply supported beam is made of a structured material, we assume that it behaves as a solid material with equivalent effective elastic properties. The product EI of the Young’s modulus and second moment of area is then replaced by the effective bending stiffness.

The product EI of the Young’s modulus and second moment of area is then replaced by the effective bending stiffness . This allows incorporating the effect of the topology and external geometry in the load-deflection behavior, since variable curvature in the lattice beams affects the shape of the external geometry.EI⟨⟩Lattice beam samples were tested in three-point bending at L = 100 mm, placing each sample in a pair of supports and aligning to apply the load at the center of the beam. For the mechanical tests, a Perten Instruments® TVT6700 texture analyzer with a 50-kilogram load cell was used. The machine was setup in displacement-control mode, with a test speed of 0.5 mm/s and a maximum displacement of 12 mm. An example of the mechanical test setup and the resulting load-displacement curve can be found on the left side of Figure 2.For each of the load-displacement curves obtained from the texture analyzer, two points were extracted, as shown on the right side of Figure 2.

In first place, the displacement δ corresponding to F = 0.5 N was extracted to calculate the bending stiffness of the lattice beam. This value was selected because it is within the linear elastic range for all the topologies tested. In second place, the peak force Fp was extracted to calculate the yield stress of the lattice beam, = 3FpL/(2d2w). Here, d is the depth of the beam in the load direction, which can be calculated as while w is the width out of the loading plane.σσ⟨⟩⟨⟩1888Figure 2. Left: Setup of the mechanical test. Right: Sample stress-strain curve corresponding to that sample. The displacement δ corresponding to F = 0.5 N, which is used for effective stiffness calculation, is marked in the plot, along with the peak force Fp used to calculate yield stress.ResultsBending stiffness results of the mechanical testing of the double curve topology are presented as a function of r/λ in Figure 3 (left side).

Values are obtained from the simply supported = FL3/(48δ), where F and δ are obtained from each stress-strain curve. beam equation, as Mean values from the three replicates were calculated and plotted with ±1 standard deviation error bars. Increasing r/λ results in a lower bending stiffness: for r/λ = 0.6 and β = 30°, the bending stiffness of the lattice beam is 72.61% smaller than for the square topology. Note that smaller values of β result in larger values of bending stiffness; for the same case with r/λ = 0.6 but with β = 90°, the resulting bending stiffness is 92.61% smaller than for the square topology.EI⟨⟩On the right side of Figure 3, bending stiffness is compared for the single and double curve topologies; the single curve topology has a larger bending stiffness than the double curve.

The bending stiffness of the double curve topology is 13.12% smaller than for the single curve topology for β = 90°, and 34.14% smaller for β = 60°.Yield stress is also calculated from the peak force and presented in Figure 4; on the left side, mean yield stress values for the double curve topology are plotted as a function of r/λ, and once again the yield stress is smaller as r/λ increases. For r/λ = 0.6 and β = 30°, the yield stress is 70.44% smaller than for the square topology. When increasing the value of β, the yield stress decreases further; for r/λ = 0.6 and β = 90°, the resulting yield stress is 94.51% smaller than for the square topology. On the right side of Figure 4, yield stress is compared between the single and double curve topologies; yield stress is also larger for this topology.

On the right side of Figure 4, yield stress is compared between the single and double curve topologies; yield stress is also larger for this topology. The yield stress of the double curve topology is 19.54% smaller than for the single curve topology for β = 30°, and 37.18% smaller for β = 60°.1889as a function of r/λ for the double curve topology; colors Figure 3. Left: Bending stiffness and markers represent different values of β. Right: Comparison of bending stiffness between single and double curve topologies (for the case with r/λ = 0.4), as a function of β. Mean values are plotted with ±1 standard deviation error bars in all cases.EIEI⟨⟩⟨⟩as a function of r/λ for the double curve topology; colors and Figure 4. Left: Yield stress markers represent different values of β. Right: Comparison of yield stress between single and double curve topologies (for the case with r/λ = 0.4), as a function of β.

Left: Yield stress markers represent different values of β. Right: Comparison of yield stress between single and double curve topologies (for the case with r/λ = 0.4), as a function of β. Mean values are plotted with ±1 standard deviation error bars in all cases.σσ⟨⟩⟨⟩Discussionof the Bézier-based lattice beams decreases as r/λ increases. This obeys the fact that a larger distance r from the center to control point P1 results in a Bézier curve with higher curvature and arclength, leading to a higher flexibility in the arrangement, especially for the double curve topology where the bending stiffness is effectively lower. Note also that increasing β also leads to an increment of the curvature, which gives lower values of β a largerThe bending stiffnessEI⟨⟩1890bending stiffness. Variability among the samples is also low, which is a measure of the repeatability of the 3D printing process.Figure 5.

Variability among the samples is also low, which is a measure of the repeatability of the 3D printing process.Figure 5. Qualitative finite element simulation to observe stress distribution in (a) the double curve topology and (b) the single curve topology, for r/λ = 0.4 and β = 90°.In the case of the yield stress, the force required to bend the lattice beam samples increases, since it is proportional to the bending stiffness. Considering that stress and force also have a linear relationship, stress increases with the bending stiffness. A qualitative finite element simulation of the lattice beams in bending was performed in terms of the stress distribution in the single and double lattice beam samples. Note from the simulation that in both the single and double topologies, the greatest stress concentration is in the horizontal members, which carry the load as opposed to the vertical members, which remain practically stress-free.

This coincides with the loading scenario, where horizontal members effectively behave as simply supported beams in bending. However, the magnitude of the peak stress is larger in the single topology. The curvature of the Bézier curved elements has an effect on the stress distribution, since peak stresses appear at the sections of the curve with the smallest minimum radius of curvature.ConclusionsTwo topologies of Bézier-based lattice beams were fabricated and tested in three-point bending to obtain the bending stiffness and peak stress of each arrangement. The bending stiffness of lattice beams decreases when increasing the distance r between the free control point and the origin. Yield stress follows a similar behavior, and the stress concentration occurs at the horizontal members transverse to the load.

Yield stress follows a similar behavior, and the stress concentration occurs at the horizontal members transverse to the load. When comparing the single and double curve topologies, the double curve arrangement has a larger flexibility and lower yield stress than its single curve counterpart.The beam length and variability of the printed samples requires the characterization of deflection due to shear as future work. Other characterization methods can be used to analyze the lattice beams presented here, in order to provide a better insight on the properties of this topology and its adaptation to suit various engineering applications. Overall, the synthesis method presented here allows the design of lattice beams with custom bending stiffness and large flexibility.AcknowledgmentsThis research was supported by Mexico’s CONACyT Scholarship Program (1006922) forPhD studies of first author. The support of Tecnologico de Monterrey is also acknowledged.1891References[1] K.

 Experimental Determination of Optimum Foil Joint Conditions for Structural Parts Fabricated by Ultrasonic Consolidation J.O. Obielodan, G.D. Janaki Ram, B.E. Stucker Mechanical and Aerospace Engineering Department, Utah State University, Logan, UT84322Reviewed, accepted September 10, 2008Abstract This paper describes an investigation of the optimum conditions necessary to eliminate defects at foil joints in parts fabricated by ultrasonic consolidation. Tensile test specimens were fabricated with different foil joint conditions of varying degrees of overlap in the deposition layers. They were subjected to tensile tests to determine their mechanical properties. Microstructures of samples were also studied. Experimental results show correlations between foil joint condition and mechanical strength. Sample microstructures also show correlations between the bonding qualities of the foil joints and the strengths obtained.

Sample microstructures also show correlations between the bonding qualities of the foil joints and the strengths obtained. The study highlights an important process parameter to control for fabrication of defect free structural members by ultrasonic consolidation.1.IntroductionUltrasonic consolidation (UC) is an innovative solid-state fabrication process that combines the ultrasonic welding of metal foils and layered manufacturing methodologies to produce solid freeform objects (Kong et al; 2004). The UC FormationTM technology developed by Solidica employs ultrasonic vibration combined with a normal force to generate static and oscillating shear forces within metal foils to produce solid-state bonds. A layer consists of one or more foils placed side-by-side such that their combined width exceeds the size of the part which is being fabricated. Foils are welded in a stack of layers to fabricate a part.

Foils are welded in a stack of layers to fabricate a part. A subtractive process is integrated in UC via a 3-dimensional computer numerical control (CNC) milling head for generating the layer by layer geometrical profile of the solid object. The geometry of each layer in the CAD file is replicated on the fabricated part by milling the foil layers to their desired contours. This additive and subtractive process fabricates near net shape parts, as only the substrate or base plate needs to be removed by milling to form the completed part. In the case where the base plate is integrated into the design, the process is a net shape process.Previous work has shown the viability of this novel additive process for fabricating various multi material solid objects including metal matrix composites (Janaki Ram et al, 2007).

Some of the applications of the UC process include fabrication of tools with conformal cooling channels, embedded electronic structures, embedded fiber optics, honeycomb structures and structures with arbitrary cavities (White, 2003; Kong et al, 2005; Siggard et al, 2006; Yang et al, 2007 ). Also, as UC is a low temperature solid freeform process, this offers a major distinction for the process when compared with other additive manufacturing processes for metals. It operates below 50% of the melting temperatures of the metals being processed. As such, thermal stresses and related problems like distortion and embrittlement in fabricated parts are minimized.Good physical and metallurgical bonding between foil layers is of paramount importance to minimizing defects in UC fabricated parts. Each material or combinations of materials have optimum parameters required to obtain structurally sound part fabrication.

Each material or combinations of materials have optimum parameters required to obtain structurally sound part fabrication. Lack of good bonds between foils at the layer interface results in low-strength parts that are unsuitable for structural applications. Also, leakage can result in cases where flow channels are incorporated in the part.Optimum process parameters have been established for the fabrication of some materials in previous studies, especially Aluminum 3003H18. Other materials for which process parameters have been established are aerospace grade aluminum such as 6061, aluminum-silicon carbide metal matrix composites, and 316L stainless steel (Kong et al, 2004; Yang et al, 2007; Tuttle, 2007). The important process parameters are vibration amplitude, normal force, welding speed, temperature, and layer surface roughness. All of these parameters have direct influence on the bond quality, and hence, the strength of the structure.

All of these parameters have direct influence on the bond quality, and hence, the strength of the structure. It has been established that linear weld density can be greatly improved by surface machining each deposited layer. Almost 100% linear weld density has been achieved by surface machining (Janaki Ram et al, 2006).Previously established process parameters largely focused on improving interlayer bonds between foils. In a typical UC fabricated part with less than 100% linear weld density, a number of interlayer bond defects, will be present. Figure 1 illustrates interlayer defects in a UC fabricated part. The process parameters mentioned above are generally optimized to eliminate or reduce the interlayer defects for a material or combinations of materials. Another area of concern for which no optimum parameter has been established is the joint between adjacent foils in each layer, that is, the edge-to-edge joint between adjacent foils within a layer.

Another area of concern for which no optimum parameter has been established is the joint between adjacent foils in each layer, that is, the edge-to-edge joint between adjacent foils within a layer. For parts with more than 24mm width (approximately one foil width); more than one foil laid side-by-side are required to create each layer. The foil joint condition in the layers, that is, whether an overlap or a gap exists, is a critical consideration. Gaps between adjacent foils are very common defects in UC deposited parts. The optimized process parameters mentioned above have not offered any solution to this problem thereby making UC fabricated parts unsuitable for some structural applications. These defects create potential stress risers that affect the load carrying capability of structures fabricated by UC. An example of such foil joint defects in an Al 3003 fabricated part is shown in the micrograph in figure 2. It reveals gaps between adjacent foils in a typical UC fabricated part.

An example of such foil joint defects in an Al 3003 fabricated part is shown in the micrograph in figure 2. It reveals gaps between adjacent foils in a typical UC fabricated part. The topmost gap is about 150μm, almost the same as the thickness of the foil. This work focuses on establishing optimum process condition to minimize defects at adjacent foil joints within a layer, thereby enhancing the strength of fabricated structures.Figure 1: Micrograph showing interlayer bonding defects150 μmFigure 2: A micrograph with arrows showing foil joint defectsIf UC fabricated parts are used in applications requiring parts to withstand dynamic loads, like in aircraft and automobiles, it becomes very important that a solution is found to foil joint defects. The weakening effects of these defects is more pronounced in complex structures that have features like ribs with transverse foil orientation to the direction of loading.

The weakening effects of these defects is more pronounced in complex structures that have features like ribs with transverse foil orientation to the direction of loading. Those features are more likely to fail early in comparison with other parts that have foils in the longitudinal direction to the applied load. A default foil width of 0.9411” (23.90mm) in the machine code for part fabrication is generally maintained in the UC machine. This width automatically sets the foil joint condition for Solidica supplied Al 3003 UC foil. In this work, tensile test specimens that model different foil joint conditions with varied tape width settings were fabricated and tested for strength. The fractured specimens were subjected to metallographic and fractographic studies to establish possible correlation between the quality of the foil joints, the mode of fracture and the strength of the specimens.

Al 3003H-18 foil (nominal composition by wt%: Al-1.2Mn-0.12Cu, 23.90mm wide and 150μm thickness) sourced from Solidica Inc., USA was used.2Experimental Work2.1FabricationThe experimental work was carried out with an ultrasonic consolidation machine built by Solidica, known as a FormationTM machine. Foil feed and positioning, welding and contour milling operations on the machine are computer numerically controlled. A machine code is generated for the CAD model of the part to be fabricated. This code incorporates the fabrication process parameters and directs the sequential operations of the UC machine. The foil positioning is generally randomized across the layers in the machine code automatically. This random arrangement almost always requires adjustments for good interlayer bonds and ease of fabrication. The machine has a heat plate, by which the substrate is regulated at the desired optimum temperature.

The machine has a heat plate, by which the substrate is regulated at the desired optimum temperature. Al 3003 base plate substrates (dimensions 355mm x 355mm x 12mm) bolted to the heat plate were used for the deposition of the foils. A rotating ultrasonic sonotrode of 150mm diameter vibrating at 20 KHz frequency travels on the automatically fed foil to weld it to the substrate. Figure 3 describes the Solidica ultrasonic consolidation machine. The following optimum process parameters already established in previous work for aluminum 3003 fabricationwere applied in the machine code for all the specimens fabricated in this work: Temperature 300oF; Normal Force: 1750N; Amplitude16μm; Welding Speed: 12mm/s.The tape width settings in the machine code were varied for each set of specimens fabricated; these width settings determine the distance the sonotrode is displaced to weld adjacent foils.

The default tape width setting in the machine code is 0.9411” (23.90mm) with standard Al 3003 foil supplied by Solidica. Since the foils are of the same width, the distance the sonotrode moves then determines the gap between adjacent foils. This gap is critical as it determines the closeness or overlap of the foils at the joints. If the sonotrode displacement exceeds the standard tape width, adjacent foils do not touch each other and a gap is created between them. However, if the displacement is less than the standard foil width, adjacent foils overlap at the joints.

However, if the displacement is less than the standard foil width, adjacent foils overlap at the joints. The extent of the gap or overlap created at the joint is dependent upon the difference between the standard foil width and the sonotrode displacement or foil width specified in the machine code.Figure 3: (a) Solidica FormationTM UC machine, (b) Close-up view of ultrasonic sonotrode from below, (c) Schematic of UC process.Figure 4: Samples of the fabricated tensile test piecesStandard tensile testing bars conforming to ASTM specifications were fabricated with the following dimensions: Gage length – 50.0mm; Width – 12.5mm; Thickness – 7.0mm; Radius of fillet – 12.5mm; Overall length – 200.0mm; Reduced section – 57.0mm; Length of grip section – 50.0mm; and Width of grip section – 20.0mm. Samples of the tensile specimens are shown in figure 4. Three tensile bars were fabricated for each foil joint model.

Samples of the tensile specimens are shown in figure 4. Three tensile bars were fabricated for each foil joint model. Solid blocks with dimensions 210mm x 75mm x 7.2mm were first deposited, from which the tensile bar profiles were machined out with the integrated CNC facilities in reverse order from the topmost layer to the first layer in contact with the substrate. The tensile bars were removed from the substrate by machining off the substrate on a conventional milling machine. Ten different sets of three tensile bars were fabricated, each modeling a specific foil joint condition2.2Longitudinal Tensile barsOut of the ten sets of specimens fabricated, one set modeled a longitudinal foil arrangement (that is, longitudinal with respect to gage length) and default tape width setting of 0.9411” (23.90mm) in the machine code. The foils were arranged with 50% overlap across layers as illustrated in figure 5a below.

The foils were arranged with 50% overlap across layers as illustrated in figure 5a below. This is like a brick wall arrangement intended to maximize the amount of overlap with previously deposited layer. The overlap arrangement is such that the joints of foils in a previous layer were located at the middle of the foils in the immediate adjoining layer.a: 50% overlapb: Random overlapFigure 5: End view patterns of 50% and random overlap foil arrangement2.3Transverse Tensile SpecimensSeven sets of bars with specific foil width settings in the machine code were fabricated to model foil arrangements in the transverse direction with respect to the gage length of the fabricated tensile test specimens. The different tape width settings for the specimens with transverse foils ranged from 0.9363” (23.78mm) to 0.9435” (23.96mm) with an increment of 0.0012” (0.03mm). These seven sets utilized 50% foil overlap between the layers, as shown in figure 5a above.

These seven sets utilized 50% foil overlap between the layers, as shown in figure 5a above. The 0.9423” (23.93mm) and 0.9435” (23.96mm) foil width settings were expected to produce gaps between adjacent foils. These excessive settings were purposefully included in order to evaluate the effects of differing amount of gaps on the strength of the modeled structures.2.4 Random Transverse Tensile BarsTwo sets of bars were fabricated to model tape width settings of 0.9435” (23.96mm) and 0.9375” (23.81mm) with random overlap. An example of the random overlap is illustrated in figure 5b. These were fabricated to compare the strengths of the random overlap and 50% overlap settings.For the purpose of identification in this work, the sets of specimens have been labeled according to their width specifications in the machine code without the “inch” and decimal notations. As an example, the 0.9363” width setting specimens are labeled 9363 specimens.

As an example, the 0.9363” width setting specimens are labeled 9363 specimens. For those with the random foil arrangements, their labeling is prefixed with an “R”, for example, the random 0.9435” specimens are labeled R9435 specimens. The longitudinal sample set is labeled “Long.”2.5 Problems EncounteredThere were problems encountered during the deposition of some of the blocks for the tensile specimens. The depositions for sample sets 9363, 9375, 9387 and 9435 were more difficult to build. The level of difficulty increased as the width difference between the default 0.9411” and the desired setting increased. Thus, 9363 sample set was more difficult than 9375 sample sets in that order; also the 9435 sample set was more difficult than the 9423 sample set. Figure 6 shows a picture of one of the deposition problems that were encountered a number of times during the fabrication of the blocks.

Figure 6 shows a picture of one of the deposition problems that were encountered a number of times during the fabrication of the blocks. Some foils popped up without bonding to the previous layer as a result of the new width settings. This problem was typically corrected by rewelding the affected foil(s). However, in some cases the entire layer was removed by a milling recovery operation. The deposition process continued from the recovered layer.Popped foilFigure 6: Picture showing a layer with a popped foil. This was corrected by repeating the weld2.6 Tensile TestingThe bars were subjected to tensile tests with an Instron tensile testing machine (model 3367) with load cell capacity of 30KN according to the ASTM standard. Also, an Instron 2630- 100 Series clip-on extensometer was used for strain measurements.

Also, an Instron 2630- 100 Series clip-on extensometer was used for strain measurements. With the extensometer, the Young’s modulus for each sample was deduced from the stress-strain relationships.2.7 Metallographic StudiesSmall samples cut from the undistorted tensile specimens’ grip ends were mounted and polished according to standard metallographic preparation procedures. They were then etched with Kellers reagent (90ml H2O, 5ml HNO3, 3ml HCl and 2ml HF) and observed under an optical microscope for microstructural characterization. Also, fractographic studies were carried out on the fractured samples with Scanning Electron Microscopy (SEM).

Also, fractographic studies were carried out on the fractured samples with Scanning Electron Microscopy (SEM). The fractured samples were ultrasonically cleaned with acetone in preparation for the SEM studies, such that dirt was removed effectively with the original fracture lines unaltered.3Results and Discussions3.1 MicrostructuresThe micrographs taken from the optical microscope reveal distinctive microstructure foreach set of specimens as shown in figure 7 below.a: A Longitudinal sampleb: A 9435 samplec: An R35 sample150 μm150 μm150 μmd: A 9423 samplee: A 9411 samplef: A 9399 sample150 μm150 μm150 μm150 μmg: A 9387 sample150 μmh: A 9375 sample150 μmi: A 9363 sampleFigure 7: Representative micrographs of each set of tensile specimensThe micrographs in figure 7 above reveal the bonding conditions at the foil joints of the respective sample sets of the specimens.

It is evident that the frequency and size of defects at the joints is proportional to the foil width specified in the machine code for the respective specimens. The 9435 sample in figure 7b, a sample of the specimens with the widest foil width setting, shows the widest gaps between the foils to the extent that a foil bent over into the gap in a lower layer can be observed. Virtually all alternate layers have sizable gaps in the sample. A trend can be observed in the size of defects from figures 7b to 7i representing a decrease in the foil widthsetting from 0.9435” (23.96mm) to 0.9363” (23.78mm). The foil joint defects progressively reduce down to the 9375 sample where the smallest defect occurrence is observed, and then slightly larger defects in the 9363 sample. It is noteworthy that the 9411 sample representing the most commonly used foil width setting for UC fabrications with the Solidica supplied Al 3003 alloys has joint defects.

It is noteworthy that the 9411 sample representing the most commonly used foil width setting for UC fabrications with the Solidica supplied Al 3003 alloys has joint defects. This can also be confirmed in the longitudinal sample represented in figure 7a that was also fabricated with a width setting of 0.9411”. It therefore means that the default foil width setting of 0.9411” for Al 3003 can not fabricate defect-free parts. It can also be observed in the micrographs that the foil edge-to-edge joints are not in perfect alignment across the layers for the 50% overlap condition, there is up to 150μm displacement between them. It shows that the UC machine does not precisely deposit foils in the location prescribed by the machine code. The precision of the machine is thus a limitation, as 150μm error is relatively high for a process that is sensitive to gap width differences of this magnitude.

The precision of the machine is thus a limitation, as 150μm error is relatively high for a process that is sensitive to gap width differences of this magnitude. From observations made on the feeding and guiding mechanism of the machine, there is more clearance than required to precisely locate the foils in the tacking positions on the substrate. Also, when the foils are not properly tacked, sometimes they slide out of position during the welding operation, thereby creating gaps between adjacent foils. In addition to these, the translation precision of the gantry on the axes of the machine can also be a contributing factor.3.2Tensile StrengthsTable 1 shows the tensile strength data obtained from the specimens. The table contains the average strength and standard deviation. Figure 8 shows the same data in table 1 in the form of a bar chart for better comparison.

Dev.St. Dev (%)7.0 6.7 6.0 11.9 2.05.42 4.20 3.81 6.64 1.08201 207 206 226192.7 201.3 197.0 233.07.2 6.0 10.1 6.63.75 2.99 5.15 2.81Strength Comparison300.0250.0) a P M200.0( h t g n e r t S e l i s n e T150.0100.050.00.0Long R9435 9435942394119399938793759363 R9375Specimen TypeFigure 8: Bar chart with error bars showing strength comparison for samples fabricated with different width settings.In the tensile test results, the 9435 samples have the lowest strength among the 50% overlap transverse tensile specimens. A trend similar to the one in the microstructure studies in 3.1 can be observed in the tensile results. From table 1 and figure 8, it can be seen that strength increases as foil width decreases. It means that as the bond quality at the foil joints improve, the strength increases.

From table 1 and figure 8, it can be seen that strength increases as foil width decreases. It means that as the bond quality at the foil joints improve, the strength increases. Since stresses are magnified at points of defects or voids, and the magnitude of amplified stress is proportional to the size of the voids, it means the increase in strength is a result of the decrease in defects at the joints. Therefore, it can be reasonably stated that the foil width specified determines the foil joint integrity, and hence the strength of a UC fabricated part. The tensile strengths peaked for the width settings of 0.9375” (23.81mm) at an average strength of 201.3 MPa as shown in table 1. The lower width value of 0.9363” (23.78mm) resulted in a smaller average strength of 197MPa. The difficulty encountered in welding the 9363 samples as a result of the reduced foil width values may have caused this reduction in strength.

The difficulty encountered in welding the 9363 samples as a result of the reduced foil width values may have caused this reduction in strength. It can be said that the 0.9375” (23.81mm) width setting is the optimum width, beyond which the part begins to degrade in strength due to an accumulation of fabrication errors. This optimum width setting is more difficult to fabricate than the default 0.9411” setting.

However, if strength is an important factor in the application of the fabricated part, it would be best to apply this optimum value.3.3Fracture features Some of the distinctive features of the fracture surfaces of the specimens as revealed bySEM are shown in figure 9.a: A view of the fractured surface of a longitudinal sampleb: Lower magnification of the longitudinal samplec: A view of the fractured surface of an R9435 sampled: A view of the fractured surface of a 9411 samplee: A view of the fractured surface of a 9387 samplef: A view of the fractured surface of a 9375 sampleFigure 9: Fractured surface of selected sample specimensa: Side view of the fracture surface of a longitudinal sampleb: Side view of the fracture surface a 9411 samplec: Side view of the fracture surface a 9387 sampled: Fracture line of a polished R9435 sampleFigure 10: Fracture lines of some of the specimensThe SEM pictures for R9435 and 9411 samples in figures 9c and 9d show the mode of fracture propagation in parts with width settings above the optimum value.

The figures show some foils with clean, unaltered surfaces at alternate layers for the 50% overlap settings and at random locations for the random overlap samples. The clean surfaces are the fracture paths through gaps created by foil joint defects, where no bonding occurred between adjacent foils. At those locations, there were no fractures in the material. The other foil surfaces with dimples are evidence of fracture at locations within continuous foils or at locations where good bonding occurred between adjoining foils.Figure 10 show the fracture lines at the side view of some of the samples. In figure 10a, there is evidence of necking before final fracture in the longitudinal samples, a characteristic of ductile failure. This is understandable since all the foils run through the length of the samples, and all of them were fractured. In contrast, all the 50% overlap samples exhibit a somewhat flat fracture surface, which is characteristic of brittle materials.

In contrast, all the 50% overlap samples exhibit a somewhat flat fracture surface, which is characteristic of brittle materials. The flatness is due to the arrangements of the foils that have the foil joint lines almost directly above one another in alternate layers. Also, because perfect bonding was not achieved in all the joints, the weakest points in the specimens were at those un-bonded and partially bonded joints. The fracture lines therefore pass through them, and in the case of 50% overlap samples, this result in a near straight fracture line. In the case of random overlap samples, represented by an R9435 sample in figure 10d, the fracture line can be seen passing through un-bonded and partially bonded foil joints.

In the case of random overlap samples, represented by an R9435 sample in figure 10d, the fracture line can be seen passing through un-bonded and partially bonded foil joints. The arrowed foils in figure 10d are some of the un-bonded foils in alternate layers, the fracture line in this case follow the random pattern of the foil arrangements.ConclusionsIn this work, it has been shown that the default foil width setting of 0.9411” (23.90mm) in UC fabrication with Solidica supplied Al 3003H-18 foil does not produce structures with optimal strength. Also, the quality of foil joints in terms of defects is directly correlated to the foil width setting in UC fabricated structures for transversely oriented foils with respect to loading direction. The higher the width set in the UC machine code, the poorer the foil joint quality for the fabricated part. The effects of the feeding and guiding mechanism for foilpositioning with respect to joint defects have been highlighted.

The effects of the feeding and guiding mechanism for foilpositioning with respect to joint defects have been highlighted. Improved feeding and guiding mechanisms will result in reduced defects, and also enhanced repeatability in foil placement.Bonding quality at the joints has been found to be directly related to the strength of fabricated structures. Data generated through tensile tests, optical metallographic and fractographic studies shows that a width setting of 0.9375” (23.81mm) for the Al 3003 standard foils yields optimal strength for UC fabricated structures with transversely oriented foils to the direction of loading.The longitudinal samples recorded better strength values than the transverse samples, in part because the foil used was manufactured by rolling, which inherently have anisotropic properties and also because of the defects at the edge-to-edge joints of the of the foils in each layer.

The drop in strength for the specimens with widths above the optimum setting of 0.9375” are considerably less than 50% of the longitudinal samples, as might be expected if no load were carried across these joints. The bonding pattern observed in the micrographs for those samples show that some of the layers have well bonded foil joints, which means the frequency of the foil edge-to-edge defects is less than the number of joints in those specimens. This reasonably explains the reason for the higher strength than expected in those set of specimens.The data generated from the randomly arranged foils i.e., R9435 and R9375 does not show general evidence of improvement over their counterparts with 50% overlap. While the R9435 show evidence of little improvement in strength over 9435 specimens, the R9375 recorded a lower average strength value when compared to 9375.

While the R9435 show evidence of little improvement in strength over 9435 specimens, the R9375 recorded a lower average strength value when compared to 9375. It therefore, can be concluded that the randomly arranged foils do not have reliable or predictable strength values when compared to the trend observed with the 50% overlap specimens.Acknowlegementthank to Science Foundation (under grants CMMI 061457 and 0522908) and Naval Research (under Grant No. N000140710633) for supporting this work.The authors wouldliketakethe opportunitytothe National the Office ofReferencesDawn R. Wright, (2003), “Ultrasonic consolidation of aluminum tooling”, Advanced materials and processes, pp. 64-65.Kong, C.Y.; Soar, R.C.; and Dickens, P.M. (2004), “Optimum process parameters for ultrasonic consolidation of 3003 aluminium.” Journal of Materials Processing Technology (v146), pp181- 187.C. Y.

PERIMENTAL DETERMINATION OF OPTIMUM PARAMETERS FOR STAINLESS STEEL 316L ANNEALED ULTRASONIC CONSOLIDATION R. Gonzalez, B. Stucker Department of Mechanical and Aerospace Engineering, Utah State University, Logan, UT 84322Abstract Ultrasonic consolidation is being investigated for building Stainless Steel structures. In this study, parameter optimization for ultrasonic consolidation of Stainless Steel 316L annealed is assessed by evaluating experimental factors of oscillation amplitude, welding speed, normal force and temperature. An L-16 Taguchi design was used to establish the statistical significance of these factors and identify the combination of processing parameters that maximizes linear welding density. Optical microscopy was performed to investigate bond quality.IntroductionUltrasonic Consolidation (UC) is an additive manufacturing process whereby layers of metal foils can be joined with a metallurgical bond by means of acoustic energy.

Moreover, the UC process has the advantage of creating metal structures without high temperatures [1]. Indeed, although localized frictional heating is involved in the UC process, the mechanism for UC is not melting [1], and thus negligible shrinkage and thermal stresses result during part building [2]. In turn, ultrasonically consolidated parts have virtually no thermal degradation in material properties. Parts may be designed and built to include complex geometric features for specific applications due to the additive manufacturing nature of the process.At its current stage of development, UC manufacturing involves ultrasonic welding and CNC milling as combined additive/subtractive manufacturing techniques for part building. As for the UC manufacturing procedures, a computer program processes a three-dimensional CAD model of the part to be built, and slices up this model into a number of horizontal layers, each layer with a thickness equal to the metal foil used.

Ultrasonically deposited foil strips are placed adjacently to each other, to create a layer. After a layer (or several layers) is completed, a computer controlled milling head shapes the layer to its slice contour. Following this, milling chips are removed and foil deposition for the next layer starts [2]. As a result of continuous addition of layers, a three-dimensional part is produced from bottom to top.A UC foil deposition schematic is shown in Figure 1. First, a thin metal foil is placed over the substrate. Following this, a rotating ultrasonic sonotrode travels along the length of the metal foil, while a normal force is applied to the metal foil through the sonotrode, keeping the foil in intimate contact with the substrate. The consolidation of the foil and the substrate is accomplished by sonotrode oscillations at an ultrasonic frequency and at user-set amplitude.

The consolidation of the foil and the substrate is accomplished by sonotrode oscillations at an ultrasonic frequency and at user-set amplitude. The direction of the sonotrode’s oscillations (direction of excitation) is along the sonotrode’s rotation axis.Figure 1. Schematic of the ultrasonic consolidation processAs a consequence of sonotrode dynamics, localized shear forces are generated from the combination of sonotrode pressure and oscillation, inducing interfacial stresses between the two mating surfaces and elastic-plastic deformation of surface asperities [1]. Furthermore, asperities deformations break up the oxide film, establishing a metallurgical bond between the foil and the substrate due to relatively clean metal-to-metal surface contact [2].

Furthermore, asperities deformations break up the oxide film, establishing a metallurgical bond between the foil and the substrate due to relatively clean metal-to-metal surface contact [2]. On a lower scale, atomic diffusion may also aid in the bonding process because local temperatures at the interface and the surrounding affected region (about 20 µm) can reach up to 50% of the melting point of the material being deposited [1]. Although still being researched, there is evidence that ultrasonic welding mechanisms for bond formation involve: i) removal of surface oxide layers, ii) plastic deformation at the interface, and iii) to a lesser extent diffusion of metal atoms across the interface.

Nonetheless, plastic deformation is considered the most important for enabling the other two mechanisms [5].There are four general control parameters within a UC system: Amplitude of Oscillation, Contact Pressure (Normal Force between the horn and foil), Welding Speed (along direction of travel) and Temperature of the base. The first three parameters depend on sonotrode interaction with the part being built. In contrast, temperature depends on the heat applied directly to the substrate, with temperature values between room temperature and 400 °F.Linear Welding Density (LWD) is the proportion of bonded area to total area within the weld interface [7]. Selection of this measure is better understood considering that ultrasonically consolidated parts typically show unbounded regions (defects, physical discontinuities) along the layer interfaces.

Selection of this measure is better understood considering that ultrasonically consolidated parts typically show unbounded regions (defects, physical discontinuities) along the layer interfaces. Indeed, the assessment of the proportional bonded region given in a LWD measurement is also important as a quality attribute for porosity in ultrasonically consolidated parts [3]. The relevance of understanding what factors influence LWD has already been observed by Janaki Ram, Yang, and Stucker (2006) in their study of UC parts. As a matter of fact, LWD strongly affects mechanical properties in the thickness direction of a UC part and the mechanical behavior of a UC structure under load-bearing stresses [4].However, optimum parameters for UC are not universal under general conditions. The magnitude of interfacial stresses at the mating surfaces during the UC process depends on currentfrictional conditions at the sonotrode/foil and substrate/foil interfaces [10] [11] [12] [13].

In general, a significant change in frictional conditions at the sonotrode/foil or substrate/foil interfaces will affect optimum UC parameter values. Therefore, it is necessary to understand that the concept of optimality of UC parameters is restricted to a reasonably consistent range of frictional conditions, whereas new optimum UC process parameters must be established if these frictional conditions change significantly.Overall, the objective of the research effort presented here is to determine the optimum processing parameters for UC of Stainless Steel 316L annealed (SS316L annealed) foils, based on maximum linear welding density criteria. SS316L was chosen due to its commercial availability in foil form, variety of applications and greater mechanical strength than Al. Additionally, research on SS316L UC will allow comparison between previous Al UC studies and this investigation.

Additionally, research on SS316L UC will allow comparison between previous Al UC studies and this investigation. Even though Kong, Soar and Dickens (2003) have recommended a joint approach for determining optimum UC parameters based on peel testing and LWD measurements, the criteria used in this study is minimizing part porosity and thus, maximum LWD is the benchmark used for determining the optimum.Literature ReviewIn previous studies, UC of 3003/6061 Al alloy structures has been investigated using welding speed, oscillation amplitude, and contact pressure as the variable process parameters [3] [6] [7]. Evaluation of the effect the aforementioned parameters had on microstructure and mechanical properties of ultrasonically consolidated parts has been the object of active research [7] [8]. In this context, selection of appropriate process parameters plays a key role in UC bond formation of Al 3003/6061 based on LWD microscopic studies and peel-off tests [7] [9].

In this context, selection of appropriate process parameters plays a key role in UC bond formation of Al 3003/6061 based on LWD microscopic studies and peel-off tests [7] [9]. Although Kong, Soar and Dickens (2003) concluded that it is possible to have a low peel load response and high linear weld density with Aluminum 6061 (due to excessive strain hardening and cycling stressing of contact points at the interface); it has been verified that a high peel load response only occurs in the presence of high LWD [9]. For instance, higher oscillation amplitude values produced higher LWD in Al 3003/6061 and either higher weld strengths (in Al 3003) [8], or no significant effect on weld strength (in Al 6061)[7]. Furthermore, Janaki Ram et al.

Furthermore, Janaki Ram et al. (2006) included substrate temperature as an additional factor for the Al 3003 UC process, and performed a comprehensive study of the effect of substrate temperature, welding speed, oscillation amplitude, surface machining and normal force on Al 3003 UC [3]. Among other things, Janaki Ram et al. (2006) concluded that higher normal force and higher oscillation amplitude increases LWD up to a certain level, beyond which LWD decreases. Additionally, it was observed that lower welding speeds (down to 12 mm/s) increase LWD, and higher temperatures produced higher LWD within a range from ambient to 350 °F [4].On the other hand, previous research has demonstrated the feasibility of the UC process using Stainless Steel 316L [14].

In addition, this study explored the role played by process parameters of welding speed, amplitude, normal force, and temperature in Stainless Steel ultrasonic consolidation, while successfully achieving initial process settings identification by peel strength testing and optical microscopy of samples. The same study presented results of ananalysis of variance (ANOVA) on SS316L UC, with peel strength as the response. Out of the four factor investigated, only amplitude and welding speed factors were statistically significant for peel strength (with a 90% confidence interval, p-value < 0.10), and amplitude exerted the strongest effect on peel strength [14].

Furthermore, as for the effect of process parameters on peel strength, higher oscillation amplitudes and lower welding speed increased peel strength; while temperature (up to 300°F) and normal force (up to 1600 N) were not statistically significant.Considering these research efforts, this study addresses parameter optimization for ultrasonic consolidation of Stainless Steel 316L annealed, by evaluating experimental factors of oscillation amplitude, welding speed, normal force and temperature in order to minimize part porosity and therefore, on the basis of maximum LWD criteria.Experimental PlanA series of experiments with SS 316L annealed (composition by weight: 16-18 %Cr, 10- 14 %Ni, 2.0-3.0 %Mo, ≤ 2 % Mn, ≤ 0.75 %Si, ≤ 0.010 %N, ≤ 0.045 %P, ≤ 0.03 %C, ≤ 0.03 %S) were performed to assess the effect of various factors in the UC process. A Solidica Formation™ machine was used to create ultrasonic consolidated samples.

A Solidica Formation™ machine was used to create ultrasonic consolidated samples. The Solidica Formation™ UC machine (Figure 2 and Figure 3) is an integrated UC building system that combines a rotating ultrasonic sonotrode, a heat plate, a foil-feeding spool mechanism, a three-axis milling head, and a software implementation for material deposition and machining [3]. Furthermore, the Solidica Formation™ sonotrode oscillates transversely according to a half-wave rectified sine wave at a frequency of 20 kHz and at user-set oscillation amplitude while traveling over the metal foil. The sonotrode itself is incorporated into a welding head and its position is controlled by numerical control.Using the Solidica Formation™ system, part fabrication is performed on a firmly bolted Al 3003 H14 (composition by weight: 0.050-0.20 %Cu, ≤ 0.70 %Fe, 1.0-1.50 %Mn, ≤ 0.60 %Si, ≤ 0.10 %Zn) base plate mounted on a heat plate.

Also, the heat plate maintains the substrate at a user-set temperature, between ambient and 400°F. A graphical description of the Al 3003 H14 base plate (dimensions: 355x355x12 mm) is shown in Figure 4, whereas plate/part fixture is presented in Figure 5.Figure 2. Solidica Formation™ machine (as shown in [3])Figure 3. Close-up view of the Welding head, showing the sonotrode from below (as shown in [3])Figure 4. Geometry of the base plate used within the Solidica Formation™ machine and bolt locations just before depositionFigure 5. Base plate and part fixture in the Solidica Formation™ machineA 146.75 mm diameter Titanium sonotrode was employed for UC depositions. Sonotrode roughness was measured using a Mitutoyo Surftest SV-602 stylus profilometer, at three evenly spaced angular locations (0°, 120°, 240°).

Sonotrode roughness was measured using a Mitutoyo Surftest SV-602 stylus profilometer, at three evenly spaced angular locations (0°, 120°, 240°). The arithmetic average of absolute values (Ra) was calculated respectively in the sonotrode direction of excitation and around the sonotrode midplane circumference (a one inch long arc centered at each angle location). Results for Ra calculations are presented in Table 1.Table 1. Schematics of the SS316L annealed UC samplesSonotrode angular locationMeasurement Type Ra (µm) in the direction of oscilations Ra (µm) around the circumference0 degrees6.065.01120 degrees4.743.96240 degrees4.816.59Furthermore a Design of Experiment (DOE) approach was chosen for evaluating UC SS316L annealed process parameters of oscillation amplitude, welding speed, normal force and temperature and performing further parameter optimization.

Four different levels for each parameter were adopted in order to evaluate any non-linear effects (Table 2). Also, specific levels for each of the parameters were selected based on Solidica Formation™ machine setting limits.Table 2. Parameters and levels for UC experimentsF ac tors T emperature C ontact Normal F orce Welding S peed Amplitude85 F 500 N 26 ipm 16 μm190 F 1000 N 38 ipm 20 μmL evels295 F 1500 N 50 ipm 24 μm400 F 1800 N 62 ipm 27 μmA special Taguchi L-16 orthogonal array, namely the L’16 (45 L-16) Taguchi orthogonal array, was utilized to determine the effects of individual process parameters. The L’16 array comprises 5 different experimental factors, 4 levels each. As only 4 process parameters are being considered in this study, the fifth factor of the L’16 orthogonal array was discarded, obtaining the experimental matrix shown in Table 3.Table 3. Taguchi L’16 experiment matrixE x periment run T emperature N. F orc e W.

Therefore, each layer comprises a single foil 63.5 mm long, 25.4 mm wide and 0.1016 mm thick, and one experimental sample resulted from each experimental run.Figure 7. Schematics of the SS316L annealed UC sampleAll 16 samples were built on a single base plate and high temperature tape was used to keep the foils in place before deposition. Completed samples on the base plate are shown in Figure 8.Figure 8. Actual experiment samples and layout of experiment runs on base plateFollowing sample deposition, a microstructural analysis of the weld interface was conducted by sectioning weld samples along the width of the foil. The samples were mounted, polished to a 0.05 µm finish and cleaned in isopropyl alcohol using an ultrasonic bath. Both interface characterization and LWD were assessed from micrograph images of samples taken from weld cross-sections.

Both interface characterization and LWD were assessed from micrograph images of samples taken from weld cross-sections. LWD was estimated by microstructural observations using a Zeiss Axiovert 100A inverted light microscope, considering only the interface between foils.As Linear Welding Density was the response measurement in this experiment, LWD wascalculated for picture frames using the following formula:Results Most parameter combinations resulted in insufficient energy to bond SS316L. In two cases significant bonding occurred but excess strain hardening energy resulted in delamination of previously formed bonds. In one case bonding was observed without delamination. In order to select successfully bonded samples a LWD threshold was set at 50%. Those samples with an average LWD greater than 50% were considered welded.

In order to select successfully bonded samples a LWD threshold was set at 50%. Those samples with an average LWD greater than 50% were considered welded. Consequently, out of 16 experiment runs performed, 3 successfully bonded samples resulted from using the process parameters combinations in the experimental matrix (Table 4). Micrographs of the 3 successfully bonded samples are shown in figures 10 through 12.Figure 9. Sample welding interfacesFigure 10. Micrograph of sample corresponding to experiment run 6Figure 11. Micrograph of sample corresponding to experiment run 15Figure 12. Micrograph of sample corresponding to experiment run 16Individual LWD measurements were carried out on successfully bonded sample interfaces.

Micrograph of sample corresponding to experiment run 16Individual LWD measurements were carried out on successfully bonded sample interfaces. The interface between the first layer and the base plate (see Figure 9) was not considered in LWD measurements because it does not constitute a Stainless Steel to Stainless Steel bond and, in contrast with the other depositions, there is no initial sonotrode-induced roughness on the substrate. Results of LWD measurements are included in Table 4.Table 4. Successful experimental runs and respective LWD resultsE xperiment run T emperature N. F orc e W.

Results of LWD measurements are included in Table 4.Table 4. Successful experimental runs and respective LWD resultsE xperiment run T emperature N. F orc e W. S peed6 15 16190 F 400 F 400 F1000 N 1500 N 1800 N26 ipm 38 ipm 26 ipmAmplitude 27 μm 27 μm 24 μm% L WD Interfac e 1 Interfac e 2 Interfac e 3 Averag e 58.22% 82.73% 52.94%90.49% 84.47% 95.96%0.54% 87.69% 62.85%83.62% 76.01% 0.00%Based on the findings presented in this work, the following was identified as the optimum process parameter combination for SS316L annealed UC: oscillation amplitude=27 µm, welding speed=38 mm/s, normal force=1500 N, substrate temperature= 400 °F.

Therefore, evidence suggests that, apart from geometry-induced effects, employing this process parameter combination should result in nearly 83% linear weld density in SS316 annealed deposits.DiscussionThe present work independently verified the feasibility of UC of SS316L annealed, obtaining 3 parameter combinations where bonding occurred. The fact that only 3 out of 16 samples had above a 50% average LWD is strong evidence that bond formation during ultrasonic consolidation of SS316L annealed is highly dependent on process parameters.Many aspects make this research effort unique. Even though SS316L UC feasibility studies have been performed before [14], LWD has not been the experimental response in anyprevious study of SS316L UC. In addition, this UC study includes specific sonotrode surface roughness data for better assessment of the existing frictional conditions.Moreover, these experimental results reveal some important trends for SS316L annealed UC.

Regarding microscopy analysis, LWD measurements across layers show well-bonded and delaminated layers for samples 6 and 16. This suggests that there is a slight degree of excess energy, and that additional experimentation may result in optimum process parameter settings which give higher LWD without delamination. A future full factorial study centered around parameter combinations similar to samples 6, 15 and 16 will be performed to more adequately search the parameter space in this region of settings.In similitude with the Al 3003 case [8], oscillation amplitude proved influential in SS316L annealed UC. Indeed, 2 of 3 successfully bonded samples occurred at 27 microns of amplitude (the highest evaluated level of this process parameter). In addition, LWD tends to increase with higher amplitudes, confirming previous results obtained with SS316L [14].Normal force values above 1000 N of force appear to contribute favorably to obtaining full width foil welds (Figure 8).

Additionally, successful depositions were obtained at welding speeds below 50 ipm, thus limiting potential build speed. With respect to substrate temperature, evidence suggests that SS316L annealed UC may be aided by higher temperatures, as samples with good bonding occurred at 190 °F and 400 °F.Unfortunately the ability to further increase oscillation amplitude and temperature is limited by the capabilities of the Solidica Formation™ machine settings. Further machine improvements that enable larger oscillation amplitudes and higher temperatures may enable more effective UC of SS316L annealed foils. Even with these machine limitations, UC of SS316L was confirmed with LWD values comparable to those found for Al 3003/6061 previously [4] [7].ConclusionThis study confirms the feasibility of the UC process with SS316L annealed foils to obtain a high degree of Linear Welding Density.

Although more experimentation is needed to obtain final conclusions and to investigate geometry-induced effects, an optimum parameter set was identified, producing approximately 83% average linear welding density by using the following combination of parameters: oscillation amplitude=27 µm, welding speed=38 mm/s, normal force=1500 N, and substrate temperature=400 °F.Acknowledgements We gratefully acknowledge financial support received from the Office of Naval Research(under Grant No. N000140710633) and financial and technical support from Solidica, Inc.References1. Daniels, H. P. C., 1965, “Ultrasonic Welding,” Ultrasonics, 3(4), pp. 190-196. 2. Yang Y., G.D. Janaki Ram and B. Stucker, 2007, “An Experimental Determination Of Optimum Processing Parameters For Al/Sic Metal Matrix Composites Made UsingUltrasonic Consolidation.”, Journal of Engineering Materials and Technology, 129, pp. 538-549.3. Janaki Ram, G. D., Yang, Y., and Stucker, B.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEXPERIMENTAL EVALUATION OF HIERARCHICAL FUNCTIONALLY GRADED LATTICES USING DIGITAL IMAGE CORRELATION AND MICRO-CTJunyang Ye*, Ata Babazadeh-Naseri*, Benjamin J. Fregly*, C. Fred Higgs III*Department of Mechanical Engineering, Rice University, Houston, TX 77005AbstractHierarchical meta-materials based on functionally-graded lattices (FGLs) have the benefit of customizable material properties. However, the effects of sharp transitions on the effective properties of FGLs have yet to be evaluated. This experimental study focused on characterizing the compressive properties of hierarchical FGLs built with smooth or sharp gradings. A total of 12 samples were 3D-printed in Ti-6Al-4V alloy and tested in axial compressive loading. Digital image correlation (DIC) was used to measure displacements and deformations.

A total of 12 samples were 3D-printed in Ti-6Al-4V alloy and tested in axial compressive loading. Digital image correlation (DIC) was used to measure displacements and deformations. The 3D-printing quality of FGLs was also evaluated by micro-CT imaging of 5 samples. The results showed that the cross-sectional areas of struts in FGLs with sharp transitions were 26% smaller than uniform lattices and FGLs with smooth transitions. Compression testing also confirmed a lower average elastic modulus in FGLs with sharp gradings.

Compression testing also confirmed a lower average elastic modulus in FGLs with sharp gradings. These results will provide insights for incorporating adjustment factors to account for the loss of strength in FGLs.Keywords: Hierarchical lattices, Manufacturing defects, Digital image correlation, Experimental, Functional graded lattice, Micro-CT, Ti-6Al-4VIntroductionRecent development in metal additive manufacturing (MAM) has extended its application to wider fields including biomedical devices for its ability to fine-tune material properties on a macro scale [1]. In particular, MAM with selective laser melting (SLM) is well suited for producing bone implants, as the printed part can mimic human bone characteristics better than solid material [2]. These human bone characteristics include graded inner structure [3] and anisotropy [4].

These human bone characteristics include graded inner structure [3] and anisotropy [4]. While 3D-printed porous metals can in theory create anisotropy, it is still hard to fine-tune the degree of anisotropy while matching bone properties. Additionally, additively manufactured implants can create rough surfaces at the bone-implant interface to facilitate bone ingrowth and fixation, producing better osseointegration results [5]. To design implants that better match natural human bones, numerous studies have tried to utilize hierarchical lattice structures made of Ti-6Al-4V (Ti64) powder because of their high freedom in tuning unit cell size and solid volume fraction (or relative density), making a variety of applications possible [6–9]. However, it is not clear how precise is MAM in producing desired material properties at the meso- scale reliably.

However, it is not clear how precise is MAM in producing desired material properties at the meso- scale reliably. Deviation in as-built products from the original design can significantly alter the effective material property, rendering one of the biggest benefits of additive manufactured implants moot. Furthermore, since SLM uses metal powder as raw material, important material properties such as the equivalent Young’s modulus compared to conventional solid material and subtractive manufacturing is still being studied [10]. Without such parameters, computational modeling will also be inaccurate in predicting the behavior of the implant.

Without such parameters, computational modeling will also be inaccurate in predicting the behavior of the implant. Therefore, it is necessary to perform experimental testing on SLM manufactured functionally graded lattices (FGL) to determine both the printing quality and product performance.Most research regarding experimentally determining the properties of hierarchical lattices involved one- directional grading, meaning that the volume fraction of the samples either simply increase or decrease along the direction of grading [12] – [16]. Few have designed samples that use more complicated grading and resemble Voigt’s iso-strain and Reuss’s iso-stress conditions, which adds more possibilities in tailoring the hierarchical building blocks to meet specific needs. Onal et al. [16] used multiple ways to weigh the graded lattice samples to calculate the as-built volume fraction.

Onal et al. [16] used multiple ways to weigh the graded lattice samples to calculate the as-built volume fraction. While weighing the samples is a popular method, it cannot give much useful information about relative printing quality when the relative density is the same but the building blocks are1894arranged differently. Scanning electron microscope (SEM) can produce high-resolution images of the surface condition of lattice samples, and has also been a popular choice in determining printing quality [15], [16], [18] – [20], but SEM can only observe the outmost struts of the lattices without knowing the conditions of internal struts.

Few studies have compared printing quality and mechanical properties for hierarchical lattice samples that have the same volume fraction but different building block combinations and therefore, different gradings.Micro Computer Tomography (micro-CT) is another popular way of examining printing quality, but most studies only analyze a selection of micro-CT slices and look for defects [18,20] only for qualitative analyses. In fact, micro-CT slices have the potential to be three dimensional reconstructed and analyzed in more quantitative methods [21]. Being able to access the 3D models of individual struts is crucial in comparing samples of the same volume fraction because weight measurements cannot provide insights in the manufacturing differences between samples if the weights do not match. It is often challenging to differentiate the differences originated from a consistent trend of misprinting certain grading patterns or simply more attached powders on lattice struts.

It is often challenging to differentiate the differences originated from a consistent trend of misprinting certain grading patterns or simply more attached powders on lattice struts. Furthermore, struts are not necessarily printed to be perfectly circular in their cross sections. In this case, only the 3D view can provide more accurate measurements of the strut cross-sectional areas.In this study we investigate and compare the mechanical properties and manufacturing-induced defects for samples of the same volume fraction and hierarchical building blocks but different gradings and loading directions through compression testing and micro-CT imaging. The functional grading used in this study is density-varying discrete unit cell grading for truss-based lattices achieved by stacking unit cells of different strut diameters together.

The functional grading used in this study is density-varying discrete unit cell grading for truss-based lattices achieved by stacking unit cells of different strut diameters together. We seek to combine the mechanical testing results and micro-CT measurements of strut cross- sectional areas to explain the effects of manufacturing process caused by the discontinuities of different severity in discrete unit cell grading on the mechanical performance of lattice structures. A total of 12 compression samples and five micro-CT samples were manufactured and tested from Ti64 powder using SLM with building blocks that have relative densities ranging between 37% and 79%. The mechanical testing results were post-processed with Digital Image Correlation (DIC) to obtain true stress-strain curves, stiffness, yield stress and yield strain.

The mechanical testing results were post-processed with Digital Image Correlation (DIC) to obtain true stress-strain curves, stiffness, yield stress and yield strain. Three types of struts, uniform struts, struts shared in smooth grading and struts shared in sharp grading, were reconstructed and measured from micro-CT samples.MethodsMechanical testingFunctionally graded lattice (FGL) structures were designed and manufactured for testing using the hierarchical design process. Body Diagonal-Face Diagonal Cubic (BD-FDC) unit cells, according to the naming convention presented by Shi and Akbarzadeh [22], were selected as building blocks for assembling 2nd order hierarchical lattice structures. Each unit cell in the assembly took one out of four possible orientations and had a strut diameter assigned from four sizes ranging from 300 μm to 600 μm that corresponded to four different solid volume fractions from 37% to 79% for each sample (Fig. 1).

1). With different combinations of a total of 16 available types of building blocks (four unit cell types, four relative densities), functional grading of the lattice structure was achieved. A more detailed description of the hierarchical design process will be communicated in a separate publication.Two sets of FGLs were investigated: lattice structures with smooth (or linear) grading and lattice structures with sharp transitions (or nonlinear grading). For smooth grading, layers of unit cells with different relative densities had maximum strut diameter differences of 100 μm, while for sharp grading, the strut diameter differences ranged between 200 and 300 μm (Fig. 2). For clarity, linearly graded lattice bodies with smooth transitions are herein referred to as L lattices whereas sharply graded lattice bodies are herein referred to as S lattices throughout the paper.

Note that the linear grading in this study also refers to stacking unit cells of discrete linear strut diameter difference, similar to some of the previous studies [23–25], while sharp grading does not follow the linear step-wise strut diameter difference pattern. All lattice structures were graded in the same y- direction, but they were tested in two directions to examine the anisotropy of the FGL. Two solid blocks were1895added to the top and bottom of the lattice structure in direction of intended loading, which could be either x or y- direction.Fig. 1. Proposed hierarchical building blocks of four unit cell orientation types and four relative densities (or solid volume fraction, VFsolid).Fig. 2. Left: Type S lattice with 200-micron and 300-micron thickness difference grading. Right: Type L lattice with 100- micron thickness difference grading.

2. Left: Type S lattice with 200-micron and 300-micron thickness difference grading. Right: Type L lattice with 100- micron thickness difference grading. Both shown in X-loading direction.Following the methodology of lattice grading described above, a total of 12 compression samples were designed and manufactured (Fig. 3). This included 3 samples in each of the four categories (x-direction type L, x-direction type S, y-direction type L, y-direction type S). The lattice bodies of the compression samples were cubic and contained 8 unit cells with a side length of 1.5 mm in each direction, thus having side lengths of 12 mm. Top and bottom square plates with side lengths of 14 mm and thicknesses of 2.5 mm were added to the compression samples to ensure even contact when testing. Top plates were engraved with batch numbers for easy identification.

Top plates were engraved with batch numbers for easy identification. Lattices loaded in x-direction were intended to resemble Voigt’s iso-strain conditions, while lattices loaded in y-direction were designed to resemble Reuss’s iso-stress condition. In each group of samples of the iso- strain or iso-stress condition, three samples had sharp grading and the other three had smooth grading. The lattice structures were designed using an implicit geometry modeling framework developed in MATLAB (MathWorks, Natick, MA, USA) and nTop (nTopology, New York City, NY, USA). The final designs were sent to 3D Systems’ on-demand manufacturing services (Rock Hill, SC, USA) for manufacturing with Ti64 alloy powder using Direct Metal Printing (DMP), an SLM process. Ti-6Al-4V Grade 23 powder was used as a raw material with a layer thickness of 30 μm.

Ti-6Al-4V Grade 23 powder was used as a raw material with a layer thickness of 30 μm. The samples were heat-treated and electrical discharge machining (EDM) processed to a surface roughness value of Ra = 2-4 μm.Compression testing was carried out on an MTS 810 machine with a maximum of 200 kips compression force and data was gathered using the MTS 793 MultiPurpose TestWare software. A metal block was placed above the samples to protect the MTS machine by preventing excessive stress that could cause the deformation of the contact surface due to the small surface area of the samples. A total of 12 compression tests were performed with one test per sample, and the loading rate was set to 0.5 mm/min in all tests. All samples were loaded monotonically until significant structural failures were observed, or displacements exceeded 6 mm. The applied force was measured via a 200-kips load cell and displacements were recorded using the internal displacement transducer of the hydraulic jack.

The applied force was measured via a 200-kips load cell and displacements were recorded using the internal displacement transducer of the hydraulic jack. The entire testing process was recorded on a cell phone camera (12 MP sensor, 77mm equivalent f/2.8-aperture). Testing of three samples were also recorded on an Olympus TG6 camera (121896MP sensor, 35mm equivalent f/2.0-aperture). Data gathered from the MTS machine was processed using an in- house MATLAB code. For each sample, the maximum slope in the elastic region was found to represent elastic stiffness. The force-displacement response of each sample was normalized, and stress-strain curve (engineering and true) were calculated. The 0.2% offset method was used to find the corresponding yield strength for each sample.Video recordings that were deemed suitable (9 of the 12) were used for Digital Image Correlation (DIC) processing to directly measure the deformation of the lattice structure at the one frame per second rate.

One video for sharply graded y-direction sample and two videos for linearly graded y-direction samples were not successful in DIC processing due to improper focus of the image. Pixel size for each footage was calibrated based on calibration patterns that were placed around the sample before starting the test. An opensource implementation of 2D Finite Element Global Digital Image Correlation (FE-DIC) in MATLAB [26] was used for DIC processing, and in-house MATLAB code was utilized for estimating the DIC-adjusted loading rate only for the lattice structure, excluding other deforming components.Fig. 3. All 12 compression samples shown. Each sample group shows the front, side, and top view. For clarity, dimensions are distributed in different photos, but all the samples are manufactured to the same size.

Each sample group shows the front, side, and top view. For clarity, dimensions are distributed in different photos, but all the samples are manufactured to the same size. a) x-direction type L, b) y-direction type L, c) x-direction type S, and d) y-direction type SMicro-CT ImagingMicro-computed Tomography (micro-CT) imaging of 3D-printed lattice structures was used for evaluating the effect of grading on printed strut sizes. Cross-sectional areas of both uniform and graded struts, located at the transition between cells with different densities, were measured from the reconstructed micro-CT slices. A Bruker SkyScan 1272 micro-CT scanner (maximum 100 μA and 100 kV) was used in this study.

A Bruker SkyScan 1272 micro-CT scanner (maximum 100 μA and 100 kV) was used in this study. Due to the limited X- ray power of the available desktop micro-CT scanner model and limited field of view, micro-CT specific samples were designed on a 4-by-4-by-4 grid using the same building block unit cells with side lengths of 1.5, making them compact enough to fit inside the field of view of the scanner. A solid plate with width of 8 mm and thicknesses of 2 mm was added at the bottom of each lattice design for marking and easy mounting onto the micro-CT platform.A total of five micro-CT samples were designed and manufactured, including four samples of uniform strut diameters ranging from 300 micron to 600 micron with an interval of 100 micron (Fig. 4a). An additional sample was designed such that it included transitions between unit cells of varying sizes, ranging from 100 microns for smoother transitions to 300 microns for sharper transitions (Fig. 4b, 4c).

4b, 4c). During scanning, the samples were placed and fixed on a rotating platform using clay. Due to the different solid volume fractions, exposure times ranged from 2500 milliseconds to 3500 milliseconds. All scans used a 0.11mm Cu filter to reduce noise. After scanning, the companion software for CT reconstruction was used to convert vertical X-ray images into horizontal slices. During this process, typical micro-CT post-processing correction methods such as alignment correction1897and ring artifacts reduction were applied. The processed horizontal slices were then rebuilt into 3-D bodies using the marching cube algorithm [27] in the form of STL files with the same pixel size as used for CT reconstruction. Geomagic Wrap 2017 (3D Systems) was then used to register the reconstructed geometries from micro-CT to the designed lattice structures, and the relevant affine transformation (i.e., rotation matrix) for perfect alignment of imaged models was extracted.

With the rotation matrices, the original 3-D bodies produced by the marching cube algorithm could be corrected to the right orientation and stored in the form of 3-D volumes in MATLAB.Fig. 4. a) Uniform strut thickness micro-CT samples from 300-600 micron with 100-micron interval. b) One side view of the graded micro-CT sample, with the center strut (marked in red) shared by the 400-micron and 500-micron side (linear grading). c) Another side view of the graded micro-CT sample, with the center strut shared by the 600-micron and 300- micron side (sharp grading).A MATLAB script was developed for counting the number of solid pixels in binarized 2-D slices by sweeping along an axis of the stored 3-D volume. Slices in the vertical direction were used for measurements due to higher quality of images in Z-direction, which was the original axis of rotation during scanning.

Slices in the vertical direction were used for measurements due to higher quality of images in Z-direction, which was the original axis of rotation during scanning. The cross- sectional area of struts were measured by finding the position of each strut and gradually expanding the upper and lower boundaries until the number of solid pixels dropped below a prescribed limit. At each expansion step, the algorithm computed the number of solid pixels that each boundary passed through within the width of the sample and compared against void pixels. To account for the surface roughness and compute the average thickness of the measured strut, each boundary stopped expanding until it passed through a target number of void pixels, which was defined as half of the theoretical number of solid pixels along that row if the sample were smooth.

In this way, the thickness of the horizontal struts could be measured for each valid 2-D slice and integrated numerically for consecutive valid slices to compute the cross-sectional area of the strut. The benefit of computing the area rather than measuring only the maximum thickness was that it could account for the irregular cross- sectional shapes of the as-built struts, not assuming that the struts were all perfectly cylindrical when comparing the design and the as-built products, thus producing potentially more accurate results.1898ResultsMechanical testing resultsForce-displacement responses of three samples of each design, as obtained from the MTS measurements, generally match well over the entire loading range for all four cases (Fig. 5).

5). It should be noted that technical difficulties with the video recording resulted in earlier stopping for tests on two samples of type L in Y-direction, thus displaying only a fraction of the entire compression testing compared to the other samples. However, since the tests interrupted after the yield point, the stiffness and yield strength could still be calculated for these samples. It is clear from the force-displacement curves that the mechanical responses of these lattice designs are different depending on the loading directions. However, the mechanical responses of sharply and smoothly graded lattices look similar in both loading directions. The processed stress-strain curves (Fig. 6) show that the overall trend in stiffness within each sample group is consistent, but some differences can still be observed. These differences may be attributed to manufacturing-induced defects which will be discussed in detail in micro-CT results.

These differences may be attributed to manufacturing-induced defects which will be discussed in detail in micro-CT results. Samples loaded in Y-direction passed yielding at lower displacements than the X-direction samples, with the stress-strain curves becoming non-linear when the strain was around 1.5%, while X-direction samples did not yield until the strain reached at least 2.5% (Table. 1). The data labeled with subscript 1 represents the material properties of the entire sample including the top and bottom plates, and the data with subscript 2 represents the material properties of only the lattice part. These two sets of data combined represent the upper and lower estimations of the stiffness based on displacements recorded by MTS.Fig. 5. Unprocessed force-displacement data directly from MTS (unit conversion from inch to mm is done).1899Fig. 6. True stress-strain curves for all samples processed from raw MTS data.

5. Unprocessed force-displacement data directly from MTS (unit conversion from inch to mm is done).1899Fig. 6. True stress-strain curves for all samples processed from raw MTS data. Solid lines represent the entire compression samples including the top and bottom plates, dashed lines represent only the lattice parts.Table. 1.

By incorporating DIC, the adjusted force-displacement data was used to calculate stiffness, yield strength and yield strain (Fig. 7). Note that since DIC only tracks the deformation of the lattice parts, DIC adjusted data only applies to the lattice part, excluding the top and bottom plates.Time (sec)Horizontal Displacement (ux)Vertical Displacement (uy)0102030Fig. 7. DIC processing snapshots for horizontal and vertical displacement in a 30-second interval (X-direction type S sample shown).1901The results show that for all samples, DIC adjusted stiffness and yield strength for porous lattice parts are larger than non-adjusted ones. Overall, the stress-strain responses computed from the DIC were comparable to those calculated from only the MTS measurements when averaging over three samples of each lattice design. The analysis of variations within each design showed that the stress-strain curves from the DIC generally would fall between those from MTS (Fig. 8).

The analysis of variations within each design showed that the stress-strain curves from the DIC generally would fall between those from MTS (Fig. 8). The standard deviation was calculated using the data points from three samples within each sample group throughout the range of DIC-processable displacement range. For the sample groups that have all three DIC processable videos, the displacement adjustment ratio from each video was applied to its corresponding sample individually. For the sample groups that have less than three suitable videos, adjustment factors for the samples without valid videos were estimated as an average from the available correction ratios within the same sample group. The average stiffness of type S samples are slightly smaller than type L samples in the same loading direction, but because of the rather substantial standard deviation, the differences are not large enough to produce statistical significance.

There are also no statistically significant differences in yield stress and yield strain between type S and type L samples in the same loading direction. Anisotropy is clearly observed, particularly in yield stress for the same grading type loaded in different directions, as the X-direction type L and type S samples have yield stresses around 290 MPa, while both Y-direction samples are around 90 MPa. Although anisotropy is less pronounced in stiffness than yield stress, statistical significance still exists to support the observation. These findings are summarized in Table. 1 and Fig 9. The data labeled with subscript DIC represents the material properties calculated from the force data given by MTS and displacement data adjusted according to DIC results.Fig. 8. Average raw true stress-strain curves and DIC processed stress-strain curves with +/- one standard deviation.1902Fig. 9. Comparison of elastic stiffness and yield strength by sample group.

8. Average raw true stress-strain curves and DIC processed stress-strain curves with +/- one standard deviation.1902Fig. 9. Comparison of elastic stiffness and yield strength by sample group. Error bars show standard deviations.Micro-CT resultsProcessing micro-CT scans revealed 25 to 40% increase in strut cross-sectional areas of as-built samples compared to original designs for uniform and type L designs, whereas the type S as-built sample only had an average of 13% increase, as is shown in Fig. 10. These differences in cross-sectional area are equivalent to 12 to 20% increase in strut widths for lattices with uniform density and type L grading, and 6% larger strut widths for sharply graded type S. The error bars represent the standard deviation and were calculated from the measurements of 9 struts per uniform design and 3 struts per graded design.

The error bars represent the standard deviation and were calculated from the measurements of 9 struts per uniform design and 3 struts per graded design. The reason why graded designs have less available struts is that the graded struts were all integrated into one sample, while the uniform designs have a total of 4 samples for different strut diameters, thus providing more data. Manufacturing-induced defects were also observed in all samples (Fig. 11). The most common defect is manufacture-induced defective pores (marked in yellow circles), which were observed in all strut thicknesses. At lower volume fractions, manufacturing-induced disconnected struts (marked in red circles) started to appear, potentially due to the strut thicknesses approaching the minimum printing feature size (254 μm).

At higher volume fractions, particularly the uniform 600-micron sample, designed void spaces (see as dark triangles in images) tended to lose geometric conformity, deforming from triangles into irregular shapes (marked in blue circles).1903Fig. 10. Comparison of micro-CT measurements and original design. U: uniform samples. A~B: the strut that is shared in grading by the side with a thickness of A and another side of B.Fig. 11. Up: reconstructed micro-CT slices for uniform samples with manufacturing-induced defects. Down: X-ray images of corresponding samples.DiscussionIn this study, a total of 12 samples were compression tested and 5 samples were micro-CT scanned. It was found that with the same relative density, linearly graded (type L) and sharply graded (type S) lattices have no statistically significant difference in mechanical properties, such as stiffness and yield stress, when loaded in the same direction.

The average value of stiffness for type S lattices were all slightly smaller than that of type L lattice, but the lack of statistical significance in the findings may be attributed to two reasons: the small sample size (n=3) and large variations within properties of 3D-printed samples from the same design, potentially due to1904random printing defects, or that the mechanical property differences caused by grading are inherently insignificant due to the printing process causing stress concentration regardless of the severity of discontinuity. The results from this study does not support the hypothesis that sharp grading can cause the material properties to decrease significantly compared to less aggressive linear grading. The compression testing did however prove that with the same building block and relative density, anisotropy could be achieved by functional grading, which was supported by significantly different yield stresses for samples loaded in different directions.

Manufacturing defects likely caused the differences within each sample group, where defective pores and disconnected struts altered the effective properties beyond the designed targets. The 300-micron struts were found to be the most vulnerable ones to become disjointed. Since all the compression samples included 300-micron building blocks, disconnected struts might have been present in most of them, causing the samples within each group to differ in mechanical behaviors.Micro-CT measurements found that the cross-sectional area of horizontal struts in uniform and type L as- built samples were systematically bigger than designs, and the percentage increase in area decreases as the strut diameter increases, indicating the possibility of a fixed increase in strut width in each layer when being printed. This could be resulted from overhang structures and attached powders, as is commonly found in SLM manufactured lattices [28].

This could be resulted from overhang structures and attached powders, as is commonly found in SLM manufactured lattices [28]. Another possibility is that since the deviation in measured diameters are in the same order of magnitude as the printing layer thickness, the minimum feature size may be approaching the minimum printing resolution of the SLM process (30 μm) that was used, causing the struts to have larger widths by about twice the printing resolution. The sharply graded struts were found to be significantly smaller in percentage increase in cross-sectional area, which may explain the smaller average values for stiffness and yield stress of type S compression testing samples compared to type L ones. Nonetheless, due to random manufacturing-induced stress concentrations, the weakest struts in the sharply graded samples may not be the graded struts.

Nonetheless, due to random manufacturing-induced stress concentrations, the weakest struts in the sharply graded samples may not be the graded struts. Thus, even if the struts were printed relatively undersized compared to the linearly graded samples, it is still reasonable for the two designs to have similar material properties, since the stiffness and yield strength are largely dependent on the weakest members.The results from mechanical testing were consistent with what was found in previous studies with similar relative density [28][11][29]. When comparing the mechanical behaviors of graded lattices, Xiao and Song concluded that step-wise graded and continuously graded lattices had similar properties [13]. Note that the grading method that Xiao and Song used was not the conventional method of adjusting strut thickness but by changing the geometry of building blocks (i.e., changing the angles between struts to make nodes closer to each other, or aspect ratio grading).

This method is less flexible in terms of tuning relative density compared to strut size-based density grading. The range of relative density achieved in their study was about 8.5% (=22.5%-14%) using rhombic dodecahedron unit cells, as opposed to 42% (=79%-37%) using the hierarchical lattice design method used in this study. Combining this study to what Xiao and Song found, and the fact that few experimental studies have achieved a statistical significance in material property differences when using different grading methods, it is speculated that when the relative density is constant, abrupt grading may not cause a significant loss of material properties, as opposed to what computational modeling suggests [30].The reason why some samples within the same sample group behave differently is not completely understood. It may have been caused by manufacturing-induced defects, as is presented in micro-CT results and is also found by Timercan, Terriault and Brailovski [20].

It may have been caused by manufacturing-induced defects, as is presented in micro-CT results and is also found by Timercan, Terriault and Brailovski [20]. Due to the size of compression samples, micro-CT scans are not possible with existing equipment because the samples exceed the dimension limits of the specimen chamber. Additionally, the cross-sectional areas of vertical and diagonal struts have not yet been studied. The SLM manufacturing process is known to produce different microstructure features induced by cooling rates during solidification when printing in different directions [31].

Finally, although the conclusion for micro-CT measurement is that sharply graded struts have a smaller increase in area compared to design than uniform and linearly graded struts, the exact reason in terms of printing parameters, laser movement and cooling is unknown, since the printer is not in-house.1905ConclusionThe main goal of this work is to experimentally test hierarchical functionally graded lattices using DigitalImage Correlation and micro-CT.Static compressive testing showed that the average elastic stiffness and yield strength of lattice structures with sharp grading (type S) is slightly smaller than type linearly graded (type L) samples, though no statistically significant results can be drawn, indicating that material properties may not have strong correlation with the severity of discontinuity in lattice structure due to discrete unit cell grading, given a fixed overall relative density.

perimental feasibility of Electrochemical Infiltration of Laser Sintered PreformsAbhishek Goel and David Bourell Laboratory for Freeform Fabrication, The University of Texas at Austin 1 University Station, MC C2200 Austin, Texas 78712 USA +512-471-3170 voice, +512-471-7681 fax agoel@mail.utexas.edu, dbourell@mail.utexas.eduAbstractThis research deals with the experimental feasibility of room temperature infiltration of Selective Laser Sintered preforms with metals. The existing principles of electrochemical deposition techniques were adapted and modified for carrying out the infiltration at low temperatures. Electroless and electrolytic deposition processes were adapted and modified to carry out metal ion infiltration and deposition within interconnected pores. The electrolytic infiltration process was modified by inserting a conductive graphite cathode in the center to draw the positive nickel ions through the interconnected porous network and to deposit them on the pore walls.

Forced diffusion method was also attempted by forcing the electrolyte through the preform at high pressures. One of the major benefits of electrochemical infiltration is low processing temperature. Low temperature reduces both energy consumption and associated carbon-footprint and also minimizes undesirable structural changes. Both conductive and non-conductive preforms may be electrochemically infiltrated, and MMCs produced by this method have potential for use in structural applications. This research is sponsored by the National Science Foundation, Grant CMMI-0926316.IntroductionThis work deals with electrochemical infiltration of parts fabricated using one of the most common powder processing additive processes, Selective Laser Sintering (SLS). Powder metallurgy and melt infiltration are the most common techniques for producing metallic matrix composites (MMCs).

Powder metallurgy and melt infiltration are the most common techniques for producing metallic matrix composites (MMCs). A common feature of these techniques is high processing temperature which is a major cause for high production costs, poor mechanical properties and high sustainability penalty. Electrochemical deposition is a technique which is most commonly used to deposit a thin film coating on various types of material substrates. In the following work the principles of electrochemical deposition are used to carry out electrochemical infiltration of metal ions into sintered preforms. One of the major advantages of electrochemical infiltration is the lower processing temperature, normally room temperature, which helps to avoid undesirable chemical and structural changes. When electrochemical techniques are applied to form MMCs, the working conditions become much more complicated.

When electrochemical techniques are applied to form MMCs, the working conditions become much more complicated. The electrolyte must transport to the interior porous structure through narrow channels and at the same time the by-products of reaction must be driven out of the preform. This work aims at facile, low-cost room-temperature infiltration of selective laser sintered preforms to create fully functional dense parts. Selectivelaser sintering is a commercial, powder-based manufacturing process that produces parts with arbitrary shape based on a computer solid model. Fig. 1 shows a schematic of the process. Powder is spread as a thin layer approximately 125 microns thick. A computer-controlled laser scans the surface and melts together the powder, creating a thin cross section of the part. Alternatively, the powder may be mixed with a transient binder which melts under the scanning laser beam and is removed or converted in a post-processing step.

Alternatively, the powder may be mixed with a transient binder which melts under the scanning laser beam and is removed or converted in a post-processing step. This process is repeated thousands of times to create an entire part with virtually unlimited geometric complexity. Indirect SLS parts are typically 50-65 percent dense with virtually all porosity being open or connected. The average pore size is typically the order of the powder size, approximately 30-60 microns. In electrochemical deposition, this pore network is filled with an electrolytic liquid, and deposition of a solid phase is initiated internally. The time required to close a pore is short since only a pore radius (~15-30 microns) needs to be deposited for closure. The porosity was measured using pychnometry and is shown in Table 1. The theoretical density of silicon carbide is about 3.22g/cc.Figure 1. Schematic of Selective Laser SinteringTable 1.

The porosity was measured using pychnometry and is shown in Table 1. The theoretical density of silicon carbide is about 3.22g/cc.Figure 1. Schematic of Selective Laser SinteringTable 1. Pychnometer Porosity Measurements of Silicon Carbide SLS Brown PartsMass (g)Apparent Volume (cc)Pychnometer Volume (cc)Apparent density (g/cc)Porosity (%)37.07 35.95 32.8525.87 25.56 23.7912.27 12.32 11.631.43 1.41 1.3852.57 51.68 51.11By control of the deposition process spatially, the electrolytic liquid may be successfully ejected in front of the deposition. First, the non-metallic/metallic phase is mixed in powder form with an appropriate binder powder. The mixture is then SLSed to create a porous “green” part bound by the binder. After this the green part is converted to a “brown” part by heating to dissociate the binder. In this way organic ligaments or bridges between individual powder particles are replaced by a residue of equal or greater bond strength.

In this way organic ligaments or bridges between individual powder particles are replaced by a residue of equal or greater bond strength. The brown part is then infiltrated by dipping it in a bath containing a metallic salt as an electrolyte. Metal ions diffuse into the porous network of the brown preform where they combine with a counter electron flowresulting in metallic deposition within the pore network [1-4]. Silicon carbide is an engineering ceramic for which the indirect SLS process has a great potential. This material is of interest where high-temperature strength and stability are required, and where wear and corrosion are issues. Some of its numerous applications include electronics processing wafer carriers and boats, kiln components and wear parts.

Some of its numerous applications include electronics processing wafer carriers and boats, kiln components and wear parts. However, silicon carbide objects, and non-metallics in general, are often difficult to manufacture due to their high melting temperatures (around 2700°C for SiC), poor sinterability, limited ductility and difficulty to machine. Selective laser sintering coupled with electrochemical infiltration offers a cost-effective and rapid manufacturing mechanism for metal-SiC composites. SiC is one of the potential candidates for high-temperature structural components. However, the most crucial aspect that limits the application of SiC-based ceramics is their brittleness. SiC is also quite difficult to densify because of the covalent nature of the Si-C bonding and a very low self-diffusion coefficient [5, 6]. Even liquid phase sintering is not able to provide highly dense SiC preforms, and the sintered preform is not free from pores.

Even liquid phase sintering is not able to provide highly dense SiC preforms, and the sintered preform is not free from pores. The average pore size in the solid phase sintered SiC preform has been seen to vary from 1µm to 100µm with an average pore size of 55µm. The porosity and strength varies according to Eq. (1):E E 0 exp(bP), 0 exp()bP(1)E,  = elastic modulus and strength of the porous structure, Eo, o = elastic modulus and strength of the non-porous structure, P = porosity and b = constant dependent on pore shape. The deposition phenomena consist of several important facets such as the electrode-solution interface, kinetics and mechanism of the deposition process which includes surface and diffusion kinetics and nucleation and growth processes. The schematic of the process is shown in Fig. 2. The mass of the deposit is w with the atomic weight of the metal Mw.

The schematic of the process is shown in Fig. 2. The mass of the deposit is w with the atomic weight of the metal Mw. The density of the deposited metal is , V is the volume of the deposited metal in cc, A is the area of the deposit in cm2 and T is the thickness in cm. The thickness of the deposit T is given by Eq. (2), where n = ionic charge and F = Faraday’s constant. w A   wM nFA   TIdt(2)For the specific case of nickel deposition discussed here, the ionic radius of nickel (+2) ions are in the range 0.065-0.073 nm, its crystal radii vary in the range 0.079-0.087 nm, and standard electrode potential for nickel E0 /Ni = -0.25V [7]. The current densities usually range from 540/600 A/m2 and the cathode efficiencies range from 93-97%. From Eq. (2), a 100 m thick layer of nickel could be deposited on a 2 cm cube in 1 hr, which is reasonable for depositing metal ions on pore walls. Xu, et al.

From Eq. (2), a 100 m thick layer of nickel could be deposited on a 2 cm cube in 1 hr, which is reasonable for depositing metal ions on pore walls. Xu, et al. showed that the deposition distribution along the depth of the pore depends on the polarization factor ξ as the ratio of interfacial reaction rate to the diffusion rate of nickel inside the porous network [10, 11]. The polarization factor is a dimensionless parameter that depends of several critical parameters such as bulk concentration of Ni2+ ions, diffusion coefficient of Ni+2 ions, overpotential and exchange current density. Hence the major factor to be taken into account is the reduction of rate of deposition on the pore surface and higher diffusion rates inside the interconnected porous network. This can be achieved by controlling the polarization factor and inhibiting the deposition on the pore surface by using additives such as sodium allylsulfonate or coumarin.

This can be achieved by controlling the polarization factor and inhibiting the deposition on the pore surface by using additives such as sodium allylsulfonate or coumarin. To increase the rate of diffusion of ions inside the porous network, external pressure may be applied to the electrolyte to force it through the preform. The pressure drop and the remaining porosity can be measured using Darcy’s law.2+NiFigure 2. Schematic of Electrochemical Deposition ProcessExperimentalTwo types of infiltration experiments were carried out: electroless infiltration and electrolytic infiltration. In preparation for the infiltration experiments, 25 mm cubic samples were made using selective laser sintering. Initial experiments were carried out for the nickel-silicon carbide system. SiC preforms were made using SLS.

Initial experiments were carried out for the nickel-silicon carbide system. SiC preforms were made using SLS. 180 grit sizes commercially available SiC powder was mixed with the phenolic binder in the ratio 10:1 placed inside ball milling jars and allowed to mix for 6-8 hours at moderate speed. The mixed batch of powder was sifted to eliminate large sized particles. Prior to that, Differential Scanning Calorimetry (DSC) was carried out on neat phenolic powder to estimate its melting and curing temperatures. The plot obtained is shown in Fig. 3. The powder was loaded into the DTM-2000 Sinterstation, and the process was carried out at a part bed temperature of 115˚C and laser power of 20W, outlines laser power of 4W and slice thickness of about 75μm. Nitrogen process gas was used to maintain the inert atmosphere inside the machine chamber.

Nitrogen process gas was used to maintain the inert atmosphere inside the machine chamber. After the green parts were obtained, these parts were burnt out in a high- temperature furnace at 180˚C/hr to 200˚C, then at 30˚C/hr to 500˚C and then at 300˚C/hr to 800˚C for brown part generation, then to 1400˚C at 250˚C/hr and finally to 1650˚C at 500˚C/hr. After this, the furnace was slowly cooled down and the final brown part was obtained.Sample Preparation: The following steps were performed in the preparation of the sample for carrying out electrochemical infiltration. Firstly, the cubic SiC preform was cleaned in acetone for 15 minutes and then washed with distilled water. The preform was sensitized in 1L aqueous solution containing 10 grams of SnCl2 and 40 ml of HCl, for 15 minutes and then again washed with distilled water. The preform was activated to open the surface pores by dipping it for 30 minutes in 1L aqueous solution 0.5 grams PdCl2 and 30 ml HCl.

The preform was activated to open the surface pores by dipping it for 30 minutes in 1L aqueous solution 0.5 grams PdCl2 and 30 ml HCl. The preform was washed and rinsed thoroughly for 10-15 minutes with distilled water [8, 9].Figure 3. DSC of Phenolic BinderElectroless Infiltration: The prepared sample was lowered into a bath containing 45 grams of nickel chloride (source), 8 grams of sodium hypophosphite (reducing agent), 100 grams sodium citrate (complexing agent) and 50 grams ammonium chloride (buffering agent) in 1L of distilled water at pH 9 and temperature of 80ºC. The bath was placed over a hot plate and was stirred using a magnetic stirrer. The process was carried out for 90 minutes [8].Electrolytic Infiltration: The process was carried out by attaching a cathode terminal of the electrode to the sample and lowering it in a bath containing 225 grams nickel sulphate, 37.5 grams nickel chloride and 30 grams boric acid in 1L distilled water at pH 3.5.

Bath temperature was maintained at 48ºC. The voltage applied between the electrodes was 2.5 V [9].Modified Electrolytic Infiltration: The electrical conductivity of silicon carbide is very low, so a sufficient electronic charge concentration was not achieved uniformly throughout the sample. This inhibited the oxidation of metal ions to form a metallic deposit. To overcome this, preforms used as the cathode were modified by inserting a conductive graphite electrode at the center of the silicon carbide preform in two configurations as shown in Fig. 4.Modified Electrolytic Infiltration of Graphite: The electrical conductivity of graphite is higher when compared to SiC. Thus similar modified infiltration technique as SiC was applied to the graphite preforms with copper rectangular cathode inserted at its center.Forced Diffusion Method: The forced diffusion method was also tried out by forcing the electrolyte through the sample using a high pressure gear pump.

The electroless forced diffusion method was tried out where no electrical connections were made to the sample. The schematic of the forced diffusion method is shown in Fig. 5(a) and the sample with the sample holder used is shown in Fig. 5(b).(a)(b)Figure 4. (a) Modified Electrolytic Infiltration: Cylindrical Graphite Cathode (b)Modified Electrolytic Infiltration: Rectangular Plate Graphite Cathode(a) (b) Figure 5. (a) Schematic of forced diffusion process (b) Graphite SampleThe advantage of this modified process is that conductive graphite will attract most of the metal ions through the pores. After the surface of these graphite electrodes are completely saturated with the nickel metal deposits, and the reaction is still continued, the nickel metal will start depositing outwards on the pores and in this way a high part density can be achieved.

After carrying out the infiltration experiments, the parts were kept inside a convection oven at 150ºC to dry them for SEM analysis and to remove trapped fluids and gases. EDS maps were obtained using a JEOL-JSM 5610 SEM equipped with an Oxford Instruments EDS system. The advantage of this modified process is postulated to be that conductive graphite will attract most of the metal ions through the pores. After the surface of these graphite electrodes are completely saturated with the nickel metal deposits, and the reaction is still continued, the nickel metal will start depositing outwards on the pores and in this way a high part density can be achieved. After carrying out the infiltration experiments, the parts were kept inside a convection oven at 150ºC to dry them for SEM analysis and to remove trapped fluids and gases.Results and DiscussionThe experiments were carried out for both electroless and electrolytic infiltration.

The EDS maps showing the distribution of nickel is shown in Fig. 6. The maps were collected at a fractured cross section near the middle of a 1 inch cube infiltrated SiC preform. Fig. 6 shows the EDS maps collected using JEOL-JSM 5610 SEM with Oxford Instruments EDS system. The results are encouraging because Fig. 6 shows that it is possible to infiltrate porous preforms with metal ions using electrochemical techniques by overcoming the surface deposition barrier. The results provide a motivation to carry out other high diffusion rate processes to completely fill the pores and obtain high density functional parts. The EDS maps show that significant amounts of nickel were seen in the preforms without any external diffusion. If an external diffusive force is applied, then higher concentrations of nickel can be seen to occupy the pores. One issue is the removal of undesirable trapped gases and fluids from the bath which may occupy a large portion of the porous network. Fig.

One issue is the removal of undesirable trapped gases and fluids from the bath which may occupy a large portion of the porous network. Fig. 7 shows the SEM images of the preform at a fractured cross-section, for the modified electrolytic infiltration – configuration b. It can be seen from these SEM images that modified electrolytic infiltration has been quite effective in depositing the metals in the pores of the SLSed silicon carbide preform. The forced diffusion method is not found to be as effective as compared to the modified electrolytic infiltration techniques. This might be due to very high pumping pressure of the electrolyte through the preform.Electroless InfiltrationElectrolytic InfiltrationModified Electrolytic Infiltration (a) Modified Electrolytic Infiltration (b)Forced Diffusion Method Modified Electrolytic Infiltration GraphiteFigure 6. EDS maps showing relative distribution of silicon, carbon and nickel after infiltrationFigure 7.

EDS maps showing relative distribution of silicon, carbon and nickel after infiltrationFigure 7. SEM images showing the fractured cross section (a) before (b) after the infiltrationTable 2 shows the weight percentage of nickel deposited in the pores inside the preform in each of the infiltration methods described above. It shows that modifying the process of electrochemical plating by inserting a conducting electrode at the center of the SLSed preform leads to an increase in the percentage of nickel deposited in the pores. Higher concentrations of nickel should be achieved by reducing the current density and applying external diffusion. The target composition of nickel to achieve 100% dense part is around 65-70% by weight.Table 2.

The target composition of nickel to achieve 100% dense part is around 65-70% by weight.Table 2. Weight % of Nickel obtained after carrying out each infiltration processProcessAverage % Nickel (by weight)Electroless Infiltration6.4Electrolytic Infiltration3.8Modified Electrolytic Infiltration 19.8Modified Electrolytic Infiltration 218.6Modified Electrolytic Infiltration Graphite16.6Forced Diffusion Method6.2It can be seen from Table 2 that the weight percentage of nickel obtained in the preform after infiltration is better in case of configuration b as compared to configuration a. Similar is the case with the modified infiltration of graphite. This is because configuration b provided a larger surface area for the ions to be deposited and especially on one side of the electrode which was parallel and facing the anode. The rate of deposition was faster on this side compared to the other side.

The rate of deposition was faster on this side compared to the other side. As a result, more amount of nickel was deposited on the walls of the pores. Hence, orientation of the two electrodes does matter in case of electrochemical deposition which can be easily resolved by using two anodes on either side of the cathodic preform. Fig. 8 shows the comparison chart for various infiltration methods and Table 3 shows the comparison between various infiltration processes in terms of energy consumption.Figure 8. Plot showing the concentration of nickel for different infiltration methodsTable 3.

Plot showing the concentration of nickel for different infiltration methodsTable 3. Energy comparison for various kinds of infiltration processProcessEstimated Power Consumption(kW)Vacuum Melt Infiltration18-24Electrochemical Infiltration0.7-0.9Forced Electroless Infiltration0.8-1.6Chemical Vapor Infiltration*6.8-10One of the major challenges faced in the current situation relates to the depletion of desired ions into the small pores (where the solution is not assumed to be well-stirred). Initial experiments were carried out just to demonstrate and observe the feasibility of the concept. While the results show incomplete closure of the void network, they are nonetheless encouraging since a significant amount of nickel was deposited deep inside the pores network.

While the results show incomplete closure of the void network, they are nonetheless encouraging since a significant amount of nickel was deposited deep inside the pores network. It is anticipated that external diffusion techniques may well force the electrolyte in and out of the porous preform to improve the extent of deposition.Conclusions and Future WorkInfiltration of SLS silicon carbide preforms with nickel led to encouraging results. The initial stages of work show that the electrochemical infiltration of SLS preforms with metals is highly feasible and has the potential to produce fully dense parts without any undesirable changes. Future work includes exploring other deposition techniques and comparison of those results to these. Ionic diffusion affects the infiltration process to a great extent and hence methods having enhanced diffusion rates such as pressurized diffusion of electrolyte into the preform need to be assessed.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceExperimental Implementation and Investigation of Real-Time Metrology for ExposureControlled Projection LithographyXiayun Zhao, Jenny M. Wang, Changxuan Zhao, Amit Jariwala, David W. Rosen*George W. Woodruff School of Mechanical EngineeringGeorgia Institute of TechnologyAtlanta, Georgia, 30332Corresponding author. Tel.: +1 404 894 9668 Email: david.rosen@me.gatech.eduAbstract Exposure Controlled Projection Lithography (ECPL) is a stereolithography based process,curing photopolymer parts on a stationary substrate. To improve the process accuracy with closed-loop control, an in-situ interferometric curing monitoring and measurement (ICM&M) system was developed to infer the output of cured height.

To improve the process accuracy with closed-loop control, an in-situ interferometric curing monitoring and measurement (ICM&M) system was developed to infer the output of cured height. The previously reported ICM&M method incorporated a sensor model and online parameter estimation algorithms based on instantaneous frequency. In this paper, to validate the ICM&M method, an application program was created in MATLAB to integrate the ECPL and ICM&M systems and to acquire and analyze interferograms online. Given the limited computing power, the interferogram analysis is performed offline. Experiments were performed curing square samples by varying exposure time and intensity. They show that the ICM&M can provide a cost-effective metrology for cured heights with excellent accuracy and reliability, and decent capability of estimating lateral dimensions. The offline ICM&M is a convincing demonstration and benchmark for the real-time ICM&M metrology.1.

The offline ICM&M is a convincing demonstration and benchmark for the real-time ICM&M metrology.1. IntroductionLack of in-situ sensors and measurement methods for in-situ process control is identified as high priority research and development activities for overcoming the barriers in improving the quality and repeatability of additive manufacturing (AM) processes [1]. Improved sensors and controls used in AM equipment and processes can enhance monitoring and control capabilities to provide real-time visibility and control of the building part. Real-time inspection and material property determination during the manufacturing process can improve production of qualified parts directly from the AM machine.

Real-time inspection and material property determination during the manufacturing process can improve production of qualified parts directly from the AM machine. Process monitors and controls should be fully integrated with the AM process.In this study, the AM process is a mask-projection stereolithography based additivemanufacturing process denoted as Exposure Controlled Projection Lithography (ECPL), which can cure photopolymer material into 3D parts on a stationary transparent substrate [2]. To improve the ECPL process accuracy, an in-situ metrology is designed to enable advanced closed- loop control. An interferometric curing monitoring (ICM) system has been developed to visualize the curing process and infer roughly the output of cured height by an implicit model and awkward phase counting which is not fast and accurate enough to measure in-process [3-5].

The author has previously reported a new interferometric curing monitoring and measurement1439(ICM&M) method distinguished by a sensor model based on instantaneous frequency and an online parameter estimation algorithm with moving horizon exponentially weighted Fourier curve fitting [6].The paper details the authors’ initial efforts to validate the ICM&M model and algorithms, asdescribed in [6]. Given the limited equipment configurations, the implementation of real-time ICM&M is mainly performed in a mimic form by playing a real-time acquired interferogram video and simultaneously performing estimation of cured height, which simulated real-time ICM&M practice is called in this research “offline ICM&M”, essentially a post-analysis of acquired video. Please note that the real-time ICM&M and offline ICM&M share exactly the same sensor model and algorithms, and the only reason we cannot achieve completely real-time ICM&M is because it demands much more computation power.

In particular, the comparison of ICM&M and microscope measured cured parts dimensions results, including both lateral width and vertical height, is presented.2. System OverviewThe overall physical system of the in-house designed additive manufacturing machine, as illustrated in Figure 1, consists of two core modules: the ECPL system (shown in the bottom blue frame) and ICM&M system (shown in the top green frame).Figure 1. Overall Physical System: the ECPL system integrated with the ICM&M systemThe ECPL system aims to deliver a serials of timed and patterned UV light beams into the resin chamber where photopolymerization occurs to form 3D object. The UV light, homogenized by the beam conditioning system and shaped by DMD, is projected through the bottom transparent substrate of the resin chamber into the photopolymer resin.The ICM&M system is based on a Mach-Zehnder interferometer [3].

A coherent laser shinesthrough a beam expander, moveable iris, and beam splitter, onto the resin chamber. Light1440reflecting off the interface surfaces of the resin chamber reflects through the beam splitter and into the camera. Due to the optical path differences between the light beams reflected from different interface surfaces, an interference pattern is observed by the camera.3. Experimental methodology3.1 ICM&M methodTo improve the process accuracy with closed-loop control for ECPL, an interferometric curing monitoring and measuring (ICM&M) method was developed addressing the sensor modeling and algorithms issues [6, 7]. In the previously reported literature, a physical sensor model for ICM&M was derived based on interference optics utilizing the concept of instantaneous frequency. The associated calibration procedure was outlined for ICM&M measurement accuracy [7].

The associated calibration procedure was outlined for ICM&M measurement accuracy [7]. To solve the sensor model, particularly in real time, an online evolutionary parameter estimation algorithm was developed adopting moving horizon exponentially weighted Fourier curve fitting and numerical integration [6, 7].Figure 2 presents a summary of the ICM&M method illustrating the relationships among theCCD camera, the acquired interferograms data, the sensor model and online algorithms. The ICM&M sensor model is developed to sense the local derivative in the interference pattern and provides a formulated problem for the parameter estimation and cured height algorithms to solve.

A time sequence of the intensity signal is needed to depict a sequence of instantaneous frequency resulting from the curing velocity, and thereby the cured height Z can be evaluated by integration as shown in the explicit sensor model.The algorithm of parameter estimation by moving horizon exponentially weighted “fourier1”curve fitting is developed to estimate the instantaneous frequency (cid:1858) in the sensor model. Note that all the blue symbols in Figure 2 denote frequency items and mappings between the model and algorithms. This procedure needs to be repeated for each subsequent time period as the part is being fabricated.

This procedure needs to be repeated for each subsequent time period as the part is being fabricated. When part fabrication is completed, an estimate of total interferogram phase angle and total cured part height are produced.To implement the measurement of cured height during on-line operation with the ECPL system, an ICM&M calibration procedure should be performed beforehand to estimate the key index of refraction difference between solid and liquid resin off-line. The overall scheme of the developed ICM&M method with evolutionary estimation and incremental accumulation enables a promising real-time implementation which will be investigated experimentally in this paper.1441Figure 2.

Scheme of ICM&M sensor method: models and algorithms [7]3.2 MATLAB application for implementing the ICM&M methodAs the ICM&M method is developed specifically for the ECPL process, to validate the method, a holistic software application is desired for implementing the ICM&M method along with the ECPL system. A MATLAB application was created to serve as a testing platform in this study as well as a prototype software for a comprehensive ECPL machine in future practice.3.2.1 Software tasksA graphical user interface using the graphical user interface development environment (GUIDE) of MATLAB was created to implement the ICM&M method for the ECPL process.

The application was designed to perform the following tasks.(cid:120) VisualizationBuild a graphical user interface to visualize the process interferograms and onlinemeasurement results.Provide user control components such as pushbuttons and edit boxes to operate the ECPL process and to streamline the process with the ICM&M acquisition and measurement analysis.(cid:120) Hardware connectivityConnect and communicate between the software application and hardware equipmentincluding the ultraviolet (UV) lamp and DMD in the ECPL system as well as the CCD camera in the ICM&M system.(cid:120) Data managementCreate a MATLAB memory map file to log the acquired interferograms data and timestamps,process parameters such as the UV lamp iris level and exposure duration, and measurement parameters such as measurement period and calibrated refractive indices.(cid:120) Numerical computation1442Implement the ICM&M algorithms, and perform data analysis required for best estimation ofthe cured height profile of the ECPL cured part.(cid:120) Report generationDocument and save into MAT-files all the data, curve fitting estimated parameters for the sensor model, along with the measurement results of the cured heights.

Display and save the results of time-height curve and ROI height profile in figures.3.2.2 Graphical user interface and manual instructionsFigure 3 shows the application’s graphical user interface (GUI). To start with the application,one should select from the left-bottom dropdown menu “Real-time Measurement” or “Offline Measurement”, for the former involves with real-time interferograms acquisition and simultaneous online measurement while the later replays off line an already acquired video of interferograms and performs ex-situ measurement. In this study, the application will perform measurement analysis while playing the frames, which could be taken as a simulation of the “Real-time ICM&M” doing acquisition and analysis simultaneously.Figure 3.

MATLAB application for the ECPL process measurement & control: GraphicalUser InterfaceMeasurement results show the measured height, estimated phase angle in unit of cycle whichis equal to 2π rad, exposed cured height and dark cured height. If multiple points are measured, the height values are average values.3.2.3 Computation environmentThe application is executable in MATLAB R2015b for 64-bit Operating System and can beused in both experiment and post analysis. The real-time acquisition of the ECPL process was done in-situ on the lab desktop computer with a processor of Intel® Core(7M) i7 CPU 870 @2.93GHz 2.94GHz and an installed memory (RAM) of 16.0 GB (8.00 GB usable). All the offline ICM&M analysis was done on an ex-situ Lenovo laptop with Intel(R) Core(TM) i7- 4510U CPU @ 2.00GHz 2.6 GHz and an installed memory (RAM) of 8.00 GB.

All the offline ICM&M analysis was done on an ex-situ Lenovo laptop with Intel(R) Core(TM) i7- 4510U CPU @ 2.00GHz 2.6 GHz and an installed memory (RAM) of 8.00 GB. Provided a more1443powerful multi-core processor and a high-speed camera, the ICM&M is expected to be able to run faster and more accurate measurement online with a full-field measurement capability if necessary.3.3 Experiment designThe MATLAB application, equipped with the ICM&M method, provides a user-friendlyplatform for streamlining the ECPL process and the ICM&M implementation as well as a powerful tool for testing the ICM&M method. According to the ICM&M sensor model and calibration process [7], the experiment design incorporates a calibration experiment and a set of validation experiment.3.3.1 Design philosophy and experiment planAs presented in the previous paper [7], a calibration procedure is required to determine the uncertain material property of refractive index in the ICM&M method.

The design of ICM&M method possesses measurement traceability by relating its measurements to a known standard – an existing well-characterized measurement system. In this study, the Olympus LEXT OLS4000 3D material confocal microscope [8] is used as a scientifically sound measurement apparatus to transfer the primary standard value of sample heights profile to the ICM&M measured object – ECPL cured part heights profile. For each new batch of material, calibration is performed once at the beginning and more times later if necessary, for instance, the material has been exposed in air too long and is likely to change in properties with oxygen diffusion etc. The derived refractive index value from calibration is assumed to be constant under the normal ECPL operating conditions so that the ICM&M measurement results are within engineering tolerance over some reasonable period of time.

In this study, a bottle of material resin was prepared and used in all the experiments which were done over a period about two weeks.To validate the feasibility and explore the capability of the developed ICM&M method, twoseries of validation experiments were designed to cure 3D square blocks for various exposure time and under different UV intensities, respectively. Each experiment category above has its own particular purpose; meanwhile together they serve to provide a thorough investigation with one common theme - to demonstrate that the ICM&M method is reproducible and robust in measuring cured parts with precision and accuracy. Table 1 presents the overall scheme of the designed experiment that consists of one calibration experiment and two validation groups. The varying process conditions for each validation group are highlighted in red.Table 1.

The varying process conditions for each validation group are highlighted in red.Table 1. Experimental Design Matrix14443.3.2 ICM&M implementationIn the implementation and investigation of the ICM&M method, there are a few practicalthoughts to address in the experiment design.First of all, choosing square blocks as target 3D objects in the experiments is because theresultant cured part with lateral aspect ratio close to if not equal 1:1 from the square bitmap projection could vividly demonstrate that ECPL process is isotropic and homogeneous and that the ICM&M method is omnidirectional.To apply the ICM&M method in an off line mode, we replayed the curing process video, andin the last frame of the interferogram that presents a better shape of the cured part we selected a line or an area of pixels as the region of interest (ROI) to measure.

Since the cured parts analyzed in this study are cuboid, theoretically it should not matter much which line to choose for measuring the height. However, considering the not-perfectly-uniform UV irradiation and stochastic process variations in the ECPL system along with some possible defects due to post- curing operation, as well as for easy visual match with the microscope measurement in later comparison of results, it is recommended to choose some features such as center lines or entire area with edge lines. The larger the ROI is chosen to be measured, the more computational time the ICM&M method needs. In the calibration, the entire area would be selected as ROI for a comprehensive analysis because its result forms the critical foundation for subsequent measurement.

In the calibration, the entire area would be selected as ROI for a comprehensive analysis because its result forms the critical foundation for subsequent measurement. In the validation experiment, to speed up the analysis without influencing the results significantly, a representative profile line would be chosen as the ROI to be measured by ICM&M.Ideally, measuring a height profile with ICM&M means measuring the cured heights of all voxels based on all pixels in the region of interest, which would be computationally tedious and costly especially in real-time ICM&M. The cured part is supposed to have a uniform heights profile due to ideally uniform UV light intensity and material properties across the curing area; hence the adjacent pixels are supposed to have very similar if not identical changed phase angles during the curing process. Interferograms pixels signal were compared, confirming the assumption of the proximity similarity in neighboring pixels’ profile.

Interferograms pixels signal were compared, confirming the assumption of the proximity similarity in neighboring pixels’ profile. It has been found that to evaluate the height profile, measuring heights for pixels in the ROI at an interval of five pixels would not affect the accuracy significantly but requires much less computation expense than measuring every pixel for the whole area. According to the recommended practice, the pixels in measured area’s lines profile is denoted in the form of Pixels (starting pixel width coordinate: interval: ending pixel width coordinate, starting pixel height coordinate: interval: ending pixel height coordinate).

For example, Pixels (245:5:365, 220) denotes a horizontal line of 25 pixels starting from Pixel (245, 220) to Pixel (365, 220) with 5 pixels between each two neighboring measured pixels.Lastly, because the UV lamp in the ECPL system is designed for users to adjust its intensityin percentage scale of the “iris level”, for example, 100% means fully open iris and maximum intensity, 0% means completely closed and no irradiation at all. Hence, we used the iris level as a nominal indication of the exposure intensity applied in the ECPL process.3.3.3 Microscope measurement practiceIn both calibration and validation experiments, we need measure the cured part using a microscope, specifically the Olympus LEXT OLS4000 3D material confocal microscope [8],1445which displays directly the profile of a selected line with height measurement result.

Since the cured part is designed to be a flat-top block, a profile line is supposed to be able to reveal the sample’s height regardless of the its position. In practice, however, cured samples are sometimes impacted by the post ECPL operations (e.g. removing and washing) and cannot present a perfectly smooth surface as will be seen in the experiment results. Hence, one should locate the to-be-measured line in a proper position where the surface is as intact as possible to obtain the actual height of the cured part. Preferably, two profile lines in X and Y directions (i.e. length and width dimensions), respectively, are measured and the average result is used as a better estimation of the actual height.As a well-characterized scientific metrology equipment, the confocal microscope provides astandard reference but could also present some measurement uncertainty due to equipment operation and sample properties.

In this study, to get as accurate result as possible, we tried to use higher magnification for all the sample measurement, and used two lines average as microscope measurement result.4. Results and Discussion4.1 Material formulationAs specified in earlier research [2], a trifunctional acrylate monomer - trimethylolpropanetriacrylate (TMPTA, SR-351) obtained from Sartomer was used as obtained, with the photoinitiator 2, 2-dimethoxy-1, 2-diphenylethan-1-one (DMPA, IRGACURE-651) obtained from Ciba Specialty Chemicals, as the resin composition for the ECPL process. The DMPA photointiator concentration in the TMPTA monomer was recommended to be 20% by weight to ensure a homogenous solution. This specific formulation required less than 30 seconds to cure a thick (hundreds of microns) layer of resin. In principle, a resin formulation with higher sensitivity could have been appropriate for ECPL.

In principle, a resin formulation with higher sensitivity could have been appropriate for ECPL. However, a fast curing resin system would impose a higher demand on faster and more accurate measurement and control; hence, the 20% resin formulation mentioned above was used. All the experiments done in this paper used the same bottle of material which consists of a 4:1 ratio by weight of TMPTA monomer (16 gram) and DMPA initiator (4 grams). The mixture was stirred for approximately 4 hours to form a homogeneous solution.4.2 Calibration experimentAccording to the experiment design presented above, for the calibration, a moderate UV intensity corresponding to the UV lamp (OmniCure® S2000) iris level at 22% was chosen so that the ECPL process cured height would not grow too fast or too slow. A 250×250 pixels square as shown in Figure 4 was displayed on DMD for 12 seconds. The ICM&M camera captured the video of interferograms when the square block part was cured by the ECPL system.

A 250×250 pixels square as shown in Figure 4 was displayed on DMD for 12 seconds. The ICM&M camera captured the video of interferograms when the square block part was cured by the ECPL system. To start offline ICM&M analysis, firstly the interferograms video was replayed in the MATLAB application, and the last frame of interferograms which shows the final cured part shape was extracted. In the video’s last interferogram as shown in Figure 5, the region of interest (ROI) was selected by human eyes recognition of the cured part outline, approximated by an area of 150×150 pixels denoted as Pixels (235: 385, 140: 290) which is formed by the four corner pixels (info of coordinates shown in the boxes) along with the outline (red dashed line).

The ICM&M method was applied to the square area of 31×31 (961) pixels designated as Pixels (235:5:385, 140:5:290), which is zoned by the corner pixels being from the top-left one in a clockwise order1446Pixel [235, 140], Pixel [385, 140], Pixel [235, 140] and Pixel [385, 140] with 5 pixels between each two neighboring measured pixels in both width and height directions. The ICM&M model and algorithm estimated the total phase angle, ∑ ((cid:1846)(cid:3036)(cid:1858)(cid:3036)) in Figure 2, for each measured pixel; and the average total phase angle is 6.150 cycles (i.e., 6.150×2π rad) as shown in Figure 6.(cid:3036)In Figure 7, two profile lines in X direction and Y direction, respectively, are measured bythe laser confocal microscope, and the average height was calculated as (cid:2875)(cid:2871).(cid:2875)(cid:2874)(cid:2871)(cid:2878)(cid:2875)(cid:2871).(cid:2870)(cid:2870)(cid:2868) 73.492 μm, which substitutes Z in the sensor model equation in Figure 2.

Calibration: outline of the cured part in the last interferogram – ROI of Pixels(235:5:385, 140:5:290) to be measuredFigure 6. Calibration: estimated individual and average total phase angles for 31×31 pixels inthe ROI of Pixels (235:5:385, 140:5:290)1447Figure 7. Calibration: measured typical profile lines in two lateral dimensions (X and Y) toestimate the height of the cured sample4.3 Validation experiment group #1: varying exposure timeAfter the calibration which resulted in a known refractive indexes difference ∆(cid:1866) = (cid:1866)(cid:3040) − (cid:1866)(cid:3039) = 0.022259, we used the same batch of resin material to further validate the ICM&M model and algorithms reported in a previous paper [7].

In the first set of validation experiments, we cured square blocks by displaying the same size (i.e., 250×250 pixels) DMD bitmap under the same UV lamp iris level of 22% as in the calibration process, but for different length of exposure time – 9 seconds (Experiment #1), 12 seconds (Experiment #2) and 15 seconds (Experiment #3), respectively.Figure 8 displays the lines of ROI for measurement in the experiments, and the datatips provide the pixels coordinates (width and height) for the end pixels of the selected line (cyan solid line in the figure) and the corner pixels of the approximated cured shape (red dashed line in the figure). As noted in Figure 8, the ending interferogram of each experiment shows a clear square shape of about 145 pixels, demonstrating that the ECPL process could preserve the cross- section shape and output a consistent size, which shape and size both can be well captured by the ICM&M system.

As in the calibration process, for each line only these pixels at an interval of every five pixels are measured to save computation time at little expense of accuracy. Hence, the measured pixels, in the selected horizontal lines in Figure 8, are actually Pixels (240:5:380, 220) in Experiment #1, Pixels (235:5:380, 200) in Experiment #2, and Pixels (250:5:395, 220) in Experiment #3.1448Figure 8. Selected line of horizontal pixels to measure with ICM&M in (a) Experiment #1: 9sexposure; (b) Experiment #2: 12s exposure; (c) Experiment #3: 15s exposureWith the chosen ROI pixels corresponding to the to-be-measured voxels, we simulated the real-time ICM&M method by replaying the video and simultaneously extracting the time series of grayscales for all the measured pixels, followed immediately by estimating the instantaneous frequency and totaling the changed phase angle which leads to the final computing of the cured heights as per the algorithms presented in literature [7].

To illustrate the details of ICM&M algorithms implementation, the sequence of figures in Figure 9 depicts the time sequence of grayscales, estimated instantaneous frequency along the time, and time curve of cured height for a typical pixel in each experiment, specifically, Pixel (220, 325) in Experiment #1, Pixel (200, 265) in Experiment #2 and Pixel (220, 280) in Experiment #3. In Figure 9(a), the grayscale has a range of [0, 255] expressing the intensity of the pixel in the interferograms captured by the CCD camera in the ICM&M system. It is not exactly sinusoidal due to the nonlinear curing process and stochastic noises including the nonlinear response of camera electronics [9]. The blue dots in the figure depict the signal data and the red line is correspondingly the fitted curve by the online parameter estimation algorithm.

The blue dots in the figure depict the signal data and the red line is correspondingly the fitted curve by the online parameter estimation algorithm. The fitted curve agrees very well with the data, demonstrating the effectiveness of the moving horizon curve fitting and capability of real-time measuring.In Figure 9(b), we estimated the instantaneous frequency consecutively every 10 frames which provided a new batch of 10 raw grayscale data points. The measurement period, e.g. 10 samples (frames) per run of measurement in this study, could be adjusted based on the computation power; and a sufficiently fast measurement is preferred to capture the process dynamic better so as to measure the process output more accurately. Figure 9(b) also shows the cumulative sum of total phase changed during the ECPL process, and the value of total phase angle substituted the item of ∑ (cid:1846)(cid:3036)(cid:1858)(cid:3036) Figure 9(c).

The total phase angle is estimated to be 5.206 cycles producing the cured height of 62.38 μm for the voxel on Pixel (220, 325) in Experiment #1. Similarly, as to the voxel on Pixel (200, 265) in Experiment #2, the total phase angle is 6.142 cycles resulting an estimated height of 73.59 μm. In Experiment #3, for the voxel on Pixel (220, 280), the total phase angle is 7.028 cycles and the estimated height is 84.21μm.

These typical voxels height estimated by the ICM&M are pretty close to the microscope measurement as will be reported in Figure 10.in Figure 2 to calculate the voxel’s cured height as shown in(cid:3036)In Figure 9(b), another thing worthy to be pointed out is that the estimated instantaneous frequencies in the three experiments are shown be alike in the first nine seconds, which makes sense because the 12-second and 15-second exposed curing experiments are actually a temporal extension of the 9-second curing under the same process conditions.The stereolithographic cure process involves mass and energy transport during the curing process, incorporating exposure and dark reaction [10], as it is vividly shown in Figure 9(a) that1449the pixel grayscales oscillation still persisted for a while after the UV light was turned off at 9, 12 and 15seconds, respectively.

As continued “dark” gelation is expected in photopolymerization [11], the cured height by the ECPL process can also be classified into exposed and dark cured height. Figure 9(c) marks the UV lamp shutdown time and displays the exposed cured height. The ECPL process is continuous while the ICM&M measurement is discrete with the digital signal of grayscales measured every 10 frames – 0.333 second with the camera acquisition speed being 30 frames per second. Hence, there is an error up to 0.333 second in the approximated UV close time.Exp. #1: exposure time 9sExp. #2: exposure time 12sExp. #3: exposure time: 15sPixel (220, 325)Pixel (200, 265)Pixel (220, 280)abcFigure 9.

#1: exposure time 9sExp. #2: exposure time 12sExp. #3: exposure time: 15sPixel (220, 325)Pixel (200, 265)Pixel (220, 280)abcFigure 9. Illustration of implementing the ICM&M algorithms in the experiments varying exposure time: (a) typical time sequence of grayscale; (b) estimated instantaneous frequency along timeline; (c) evolving cured height of the voxel on the selected the pixel.As explained in the design of experiment, a single pixel measurement could not be sufficientor conclusive. More pixels need to be measured for final height estimation with less bias. We carried on the same procedure illustrated above to calculate the cured height for all the other voxels on the selected line of pixels in Figure 8, evaluated the average height as the final result, and checked its accuracy by comparing with the microscope measurement.

As a result, the line1450height profile measured by ICM&M is represented by 30 pixels in all three experiments with an estimated average as shown in the left column in Figure 10.ICM&MMicroscopeMicroscopeEstimated Profile Line(Horizontal Profile Line)(Vertical Profile Line)(a)(b)(c)Figure 10. Line heights profile: ICM&M result VS. microscope measurement for (a)Experiment #1: 9s exposure; (b) Experiment #2: 12s exposure; (c) Experiment #3: 15s exposureThe cured parts in the experiments were measured with the Olympus 3D confocal microscope as shown in the middle and right columns in Figure 10. The “Height” value displayed at the right-bottom box of the confocal microscope screenshot in Figure 10 indicates the measured height for the selected profile line.

The “Height” value displayed at the right-bottom box of the confocal microscope screenshot in Figure 10 indicates the measured height for the selected profile line. As pointed out in the experiment design, the1451average of measured heights for two lines in X-direction (horizontal) and Y-direction (vertical) is used as the final microscope measured height result to reduce the sample measurement bias. The ICM&M resultant average height is in good accordance with the microscope measured height in these experiments.Table 2 compares quantitatively the measurement results in terms of absolute error and relative error, which were calculated assuming that the microscope measured two profile lines average was the actual value.

The agreement between the ICM&M method and the confocal laser microscope measurements demonstrates that the ICM&M method is capable of measuring the vertical height for ECPL cured parts with discernment for micron order difference and measurement range at least of about one hundred microns.Table 2. Measurement results of validation experiments varying exposure timeExposure TimeROI Pixels (Width, Height) Total Phase (cycle, i.e.

The procedure of ICM&M implementation and microscope measurement was the same as presented in the previous section about the experiments varying exposure time.For each intensity level, we cured three samples and measured each sample with the offlineICM&M module in the developed MATLAB application and the confocal microscope. To demonstrate the ECPL process difference under different intensities and how the ICM&M method respond to the process changes, we presented a representative pixel’s interferogram grayscale signal and the ICM&M method’s evolutionary estimation of the instantaneous frequency and cured height, as shown in Figure 11.In Figure 11 (a), the grayscale signals vividly reflect the process dynamics difference due to the UV intensity variations. Conforming to the photopolymerization mechanism, the higher UV light intensity is provided, the more photo initiators are in the process, and the quicker the curing is.

Conforming to the photopolymerization mechanism, the higher UV light intensity is provided, the more photo initiators are in the process, and the quicker the curing is. Within the same exposure time of 12 second, there was almost one more cycle in 35%-iris-1452level UV curing than in the 25%-iris-level curing, which furtherly had about half cycle more than the 15%-iris-level UV curing did.The process dynamics change caused by varying intensity is further confirmed in Figure 11 (b), which shows that the estimated instantaneous frequencies is larger for higher UV intensity induced ECPL process.Exp. #1: UV iris 35%Exp. #2: UV iris 25%Exp. #3: UV iris 15%abcFigure 11.

#1: UV iris 35%Exp. #2: UV iris 25%Exp. #3: UV iris 15%abcFigure 11. Illustration of implementing the ICM&M algorithms in the experiments varyingexposure intensity: (a) typical time sequence of grayscale; (b) estimated instantaneous frequency along timeline; (c) evolving cured height of the voxel on the selected the pixel.Figure 12 exemplifies the heights profile measurement results. In Figure 12(a), a selected lineis shown in the ending interferogram of a sample in each experiment. Figure 12(b) displays the ICM&M estimated heights for the pixels in the ROI lines at an interval of five pixels, and similar variations present in the heights profile as seen and explained in the previous validation experiment group.

Figure 12(c) and Figure 12(d) show the measured profile lines in X-direction and Y-direction, respectively, in order to demonstrate that the cured part indeed has a height distribution with some variation of a few microns; hence the average of two profile lines height1453values is used as an estimation of the actual cured height to assess the ICM&M results as shown in Table 3.Table 3 summarizes the measurement results in terms of absolute error and relative error,which compared the ICM&M estimated height (shaded in green) with the average of the microscope measured profile lines heights (shaded in grey). In this set of experiment, the absolute errors are less than 2 μm, which are attributed to the cued part surface roughness and profile normal distribution with expected deviations; and the less than 2% relative errors further demonstrate that the ICM&M method has excellent accuracy and sensitivity to the process input.Exp. #1: UV iris 35%Exp. #2: UV iris 25%Exp.

2π rad) 7.365 6.466 4.924Cured Height (μm)88.25 77.48 59.01Microscope Measured Profile Line Height (μm) X- Y- direction direction 90.060 88.997 77.905 78.499 58.985 60.006Average89.53 78.20 59.50Absolute Error (μm)1.28 -0.72 -0.494.5 Lateral measurement potentialityThough we focused on measuring the cured height in this study, a byproduct of lateral dimensions measurement is also available. In Section 4.3, the resultant interferograms in all the three experiments show similar size of cured shape, which have approximately 145 pixels in the line corresponding to the width as shown in Figure 10. All three experiments used the same size bitmap as DMD pattern (250-by-250 pixels), and ICM&M method could retain that information of size similarity.

All three experiments used the same size bitmap as DMD pattern (250-by-250 pixels), and ICM&M method could retain that information of size similarity. Other than being able to qualitatively measure the lateral shape, the ICM&M is found to be able to measure quantitatively lateral dimensions according to the following calculations, which estimated successfully the width of the cured squares in Figure 10.In an interferogram acquired by the MATLAB application, one pixel is actually a binning of four original pixels captured by the ICM&M camera. The camera pixel size is 2.2um-by-2.2um, hence a binning pixel corresponding to a pixel in the interferogram (e.g. shown in Figure 8) is 8.8um-by-8.8um. Because the ICM&M optics adopts a vertical beam path reflected upward from the sample to the camera, in principal it should be a 1:1 ratio mapping.

shown in Figure 8) is 8.8um-by-8.8um. Because the ICM&M optics adopts a vertical beam path reflected upward from the sample to the camera, in principal it should be a 1:1 ratio mapping. The ICM&M measured lateral size could be estimated by multiplying the value of 8.8 μm with the number of pixels in the width dimension in the cured area shown in the interferogram (Figure 8). Hence, the estimated width of the cured square part is calculated as 145 × 8.8 = 1276(cid:2020)(cid:1865), which is in good accordance with the microscope measurement results of width - 1278.6 μm on average of all the six profile lines width as shown in Figure 10.1455Relative Error1.43% -0.93% -0.82%Conclusively, the ICM&M system has great potentiality in measuring the lateral dimesnions of width and length as well as the vertical heights. We recommend in future work utilizing image analysis with image edge detection algorithms to measure the pixels length of cured shape for lateral dimensions information.

We recommend in future work utilizing image analysis with image edge detection algorithms to measure the pixels length of cured shape for lateral dimensions information. This approach of measuring lateral dimensions for cured part is actually an alternative to the current plan of measuring the shape by measuring the heights profile across the curing area, which comes back to our research scope and priority in this study.5. ConclusionIn this paper, to validate and fulfill the developed ICM method, an application program wasdesigned and created in MATLAB Graphical User Interface Development Environtment (GUIDE).

ConclusionIn this paper, to validate and fulfill the developed ICM method, an application program wasdesigned and created in MATLAB Graphical User Interface Development Environtment (GUIDE). The app was deployed onto the physical system integrating the ECPL and ICM(cid:712) automating the ECPL process by controlling the ultraviolet lamp and digital micromirror device (DMD) mask display as well as synchronizing the ICM&M system which acquires interferograms online and can analyze the acquired data online also as long as sufficient computation power is available. In this study, given the limited equipment configurations and computation resource, the data analysis and measurement computation for ICM&M was performed off line.

In this study, given the limited equipment configurations and computation resource, the data analysis and measurement computation for ICM&M was performed off line. A coherent implementation of the ECPL process and ICM&M measurement was conducted with a series of experiment curing square samples by varying the factors of exposure time and intensity, to verify the measurement characteristics including discriminability, sensitivity, accuracy, precision and repeatability. A representative profile line was measured for each cured sample by both the in-house ICM&M and a commercial confocal microscope, which by comparison are in good accordance.Future work includes curing more samples and measuring entire cured part with the ICM&Mmethod to further verify its global measurement capability and repeatability.

Also a more thorough investigation of the ICM&M characteristics including traceability, sensitivity, accuracy(cid:712) precision and dispersion will be conducted.AcknowledgmentThis material is based upon work supported by the National Science Foundation under Grant No. CMMI-1234561. Any opinions, findings, and conclusions or recommendations expressed in this publication are those of the authors and do not necessarily reflect the views of the National Science Foundation. All the related research is patent pending.Reference[1] Measurement Science Roadmap for Metal-Based Additive Manufacturing. 2013, National Institute of Standards and Technology (NIST) [2] Jariwala, A.S., Modeling And Process Planning For Exposure Controlled Projection Lithography. Ph.D. thesis, Mechanical Engineering, Georgia Institute of Technology, Atlanta, USA, 2013. [3] Jariwala, A.S., R.E. Schwerzel, and D.W. Rosen, Real-Time Interferometric Monitoring System For Exposure Controlled Projection Lithography.

PERIMENTAL INVESTIGATION OF DIFFERENT CELLULAR LATTICE STRUCTURES MANUFACTURED BY FUSED DEPOSITION MODELINGOsman Iyibilgin, Cemil YigitDepartment of Mechanical Engineering, Sakarya University, TurkeyMing C. LeuDept. of Mech. & Aerospace Eng., Missouri University of Science & Technology, USAAbstractExperimental tests were conducted to evaluate the compressive properties (yield strength and compressive modulus) and build time for five different cellular lattice structures fabticated by the Fused Deposition Modeling (FDM) process. The lattice structures had repeating unit cells, and the shapes of the unit cell under study included honeycomb, square, diamond, triangle, and circle. Test specimens were manufactured by a Stratasys Fortus 400mc machine using ABS (Acrylonitrile Butadiene Styrene) as the part material.

Test specimens were manufactured by a Stratasys Fortus 400mc machine using ABS (Acrylonitrile Butadiene Styrene) as the part material. The five different lattice structures were compared with each other and also with the sparse and sparse-double dense build styles that are directly available from the Fortus machine. Honeycomb structure was found to have the best compression properties for the same porosity, although the differences among the different lattice structures were small (<7%). All of these lattice structures were found to have much higher strength than the specimens with the same porosity built using the sparse and sparse- double dense styles. However, the various lattice structures required significantly longer build times than the sparse and sparse-double dense builds. For the honeycomb structure, our investigation also included the effects of porosity and cell size. Higher porosity led to lower compression strength but shorted build time.

For the honeycomb structure, our investigation also included the effects of porosity and cell size. Higher porosity led to lower compression strength but shorted build time. For the same porosity, the yield strength could be increased and the build time shortened simultaneously by having a certain cell size.1. IntroductionThe Fused Deposition Modeling (FDM) process is an addictive manufacturing (AM) technology that builds a three-dimensional physical part by using a work head to melt a thermoplastic supplied in the form of a wire or filament and to extrude the molten thermoplastic through a small nozzle to deposit the material along a pre-planned path for each cross-section of the CAD model of a three-dimensional part [1,2].

By building a part layer-by-layer, this technology allows for manufacture of parts with complex shapes including internal cellular lattice structures [3], which are very desirable for applications that require lower weight yet with sufficient strength such as in aerospace components [4, 5, 6].The FDM process is a cost-effective AM process that offers great ease and flexibility in fabricating thermoplastic parts [7, 8]. A previous study showed that the yield strength and stiffness of an FDM part is inversely proportional to porosity [9]. Fang et al. [10] observed that Young’s modulus tends to decrease with increase in porosity.In the present study, five different cellular lattice structures with the same porosity were produced by the FDM process to evaluate the performance of the different lattice structures compared with each other and also with the sparse and sparse-double dense build styles that are readily available from the Fortus machine for fabricating sparse parts.

The comparisons were done experimentally and they included compressive properties (yield strength and compressive modulus) and build time. The effects of porosity and cell size were also investigated for the honeycomb structure.2. Experiment2.1. Fused Deposition Modeling of Test PartsThe basic materials that can be used by the Fortus machine to build parts include ABS (Acrylonitrile Butadiene Styrene), ULTEM, PC, and PPSF [11-13]. The lattice structures in our study were fabricated with ABS-M30 material, which is 25-70% stronger than standard ABS.Our FDM process was carried out by Stratasys Fortus 400mc machine. This machine has the accuracy of ± .127 mm - .0015 mm per mm and the maximum build dimensions of 406 x 356 x 406 mm3. The layer thickness depends on the exit nozzle diameter, which ranges from 0.330 to 0.127 mm for the Fortus 400mc machine.

The layer thickness depends on the exit nozzle diameter, which ranges from 0.330 to 0.127 mm for the Fortus 400mc machine. In our study, the nozzle type was T16 (254 µm in diameter), and the layer thickness was 0.254 mm.The fabricated test parts were cylindrical with 3.81 cm in diameter and 2.54 cm in height. The shapes of the cellular lattice structures included honeycomb, square, diamond, circle, and triangle, as shown in Fig. 1. Also shown in this figure are parts produced by the Fortus 400mc machine using the sparse and sparse-double dense build styles, in order to compare their performance with the different cellular lattice structures on the compressive properties and build time for test specimens having the same porosity.

The sparse parts were built with 5 contours, 45o-45o raster angle, and 0.020 cm in raster air gap, and the sparse-double dense parts were built with 3 contours, 45o-45o raster angle, and 0.036 cm in raster air gap.Since the honeycomb lattice structure was found to have higher compressive strength than the other lattice structures, additional test specimens of the honeycomb structure were built for further compression tests in order to investigate the effects of porosity and cell size on compressive properties and build time. Variations in cell size including edge length and edge width for the test specimens with the honeycomb lattice structure are shown in Fig. 2. Figure 2(a) shows variations in edge length while keeping the edge width constant, thus the porosity increases with increase in edge length. Figure 2(b) shows variations in edge length while having the edge width changed correspondingly to keep the porosity constant.a) b) c) d) e)f) g)Figure 1.

Figure 2(b) shows variations in edge length while having the edge width changed correspondingly to keep the porosity constant.a) b) c) d) e)f) g)Figure 1. Different cellular lattice structures: a) Honeycomb, b) Square, c) Diamond, d) Circle, and e) Triangle; f) Sparse, g) Sparse-Double Dense.a)b)Figure 2. Variations in cell size for the honeycomb lattice structure: a) different edge length and same edge width, and b) varying both edge length and edge width to have the same porosity.2.2 Compression TestsCompression tests were performed using an INSTRON 4469 machine, which was equipped with a load cell having its measurement range between -50 and +50 kN. A photo of the experimental apparatus is shown in Fig. 3.

A photo of the experimental apparatus is shown in Fig. 3. The INSTRON machine is controlled by Bluehill 2 software that allows adjustment of the test parameters using a closed-loop digital data acquisition unit.The speed of 0.51 cm/min was used in the compression test, and the recorded data included yield strength and compressive modulus. The data were collected and averaged over five specimens for each set of parameters.Load CellData Acquisition UnitCompression ApparatusFigure 3. Photo of the experimental apparatus for the compression tests.3. Results and Discussion3.1.

Photo of the experimental apparatus for the compression tests.3. Results and Discussion3.1. Effects of cellular lattice structureThe effects of cellular lattice structure on the compressive properties and build time were compared among the specimens having honeycomb, square, diamond, circle, and triangle lattice structures, as well as those built with the sparse and sparse-double dense build styles available directly from the Fortus machine.The experimental results obtained were shown in Fig. 4 for the yield strength and in Fig. 5 for the compressive modulus. All of the specimens in these comparisons had the same porosity. The data obtained from the compression tests indicate that the honeycomb cellular lattice structure has the highest yield strength and compressive modulus, and all of the cellular lattice structures have much higher yield strength and compressive modulus than the sparse and sparse-double dense build styles.

The yield strength of the honeycomb structure is 217% higher than the sparse- double dense and 253% higher than the sparse build. The compressive modulus of the honeycomb structure is 286% higher than the sparse-double dense build and 579% higher than the sparse build.The stress-strain curves obtained from the compression tests on the various cellular lattice structures and also the sparse and sparse-double dense builds are shown in Fig. 6. All of the stress-strain curves have similar trends in the elastic region for the various lattice structures. However, the stress-strain curves for the sparse and sparse-double dense builds have somewhat different trends. The maximum strength obtained was between 3.8% and 4.0% for the various lattice structures, 6.8% for the sparse-double dense build, and 24.6% for the sparse build. The parts fabricated using the sparse build appear to be more elastic than those using the sparse- double dense build and those with cellular lattice structures.

The parts fabricated using the sparse build appear to be more elastic than those using the sparse- double dense build and those with cellular lattice structures. Figure 7 provides the details of stress-strain curves for the various lattice structures for the range of 3% to 6% strain.40.036.036.7135.8935.9236.0535.22) a P M32.028.0( h t g n e r t S d l e i Y24.020.016.012.016.9614.528.04.00.0Figure 4. Yield strengths obtained experimentally for different lattice structures and the sparse and sparse-double dense builds at the same porosity.) a P M1,6001,4001426.3 1425.01351.11418.71333.2( s u l u d o M e v i s s e r p m o C1,2001,000800600400200498.9246.30Figure 5.

Compressive moduli obtained experimentally for different lattice structures and the sparse and sparse-double dense builds at the same porosity.504540Hexagon) a P M(s s e r t S3530252015SquareDiamondCircleTriangle10S&D Dense5Sparse0036912 15 18 21 24 27 30Strain (%)Figure 6. Stress-strain curves for different cellular lattice structures at the same porosity.) a P M(s s e r t S45 44 43 42 41 40 39 38 37 36 35HexagonSquareDiamondCircleTriangle3.03.54.04.55.05.56.0Strain (%)Figure 7. Stress-strain curves in the range of 3% to 6% strain for different cellular lattice structures at the same porosityThe build times for the various cellular lattice structures and the sparse and sparse-double dense builds at the same part porosity are shown in Fig. 8.

8. The build times of the specimens for the various lattice structures ranged between 40 and 47 minutes, while the build times for the sparse and sparse-double dense parts were 20 and 21 minutes, respectively. The build times for the cellular lattice structure were approximately twice as much as the build times for the sparse and sparse-double dense parts.n i m e m T d l i u B,i50 45 40 35 30 25 20 15 10 5 040443940472120Figure 8. Comparison of build times for different cellular lattice structures and the sparse and the sparse & double dense builds at the same porosity.3.2. Effects of porosityAdditional test specimens of the honeycomb structure were built to investigate the effects of porosity on compressive properties and build time. Figures 9 and 10 show the changes in the yield strength and compressive modulus, respectively, vs. porosity for the honeycomb cellular lattice structure.

Figures 9 and 10 show the changes in the yield strength and compressive modulus, respectively, vs. porosity for the honeycomb cellular lattice structure. The data indicate that both yield strength and compressive modulus decrease with increase in porosity. This expected because increase in part porosity reduces the amount of material in the part. In comparison, the honeycomb structure having 57% in porosity is 36.1% higher in yield strength and 29.6% higher in compressive modulus than the honeycomb structure having 71% in porosity.) a P M( h t g n e r t Sd l e i Y50 45 40 35 30 25 20 15 10 5 036.8628.9526.6123.54556065 Porosity %7075Figure 9. Yield strength vs. porosity for the honeycomb structure1,700a P M1,600, s u l u d o M e v i s s e r p m o C1,5001,4001,3001,2001,1001529.11292.01218.81077.21,0005560657075Porosity, %Figure 10. Compressive modulus vs.

Compressive modulus vs. porosity for the honeycomb structureThe stress-strain curves obtained from the compression tests for the honeycomb specimens with different porosities are shown in Fig. 11. The stress-strain curves all have linear relationships, showing an elastic behavior in the beginning of loading. The maximum strength was obtained at approximately 3.5% strain for all the specimens with different porosities.a P M, s s e r t S50 45 40 35 30 25 20 15 10 5 056% porosity 64% porosity 66% porosity 71% porosity0246810 12 14 16 18 20Strain, %Figure 11. Stress-strain curves for the honeycomb structure at different porositiesThe build times for the honeycomb specimens with different porosities are given in Fig. 12, which shows that build time decreases with increase in porosity as expected.

12, which shows that build time decreases with increase in porosity as expected. The build time of the honeycomb structure having 57% in porosity is 85.3% higher than the honeycomb structure having 71% in porosity.70n i m e m T d l i u B,i605040302063.040.035.034.0100 55.0060.0065.0070.0075.00Porosity, %Figure 12. Build time vs. densities for the honeycomb structure at different porosities3.3. Effects of cell sizeHoneycomb specimens having different cell sizes with the edge length ranging from 0.31 cm to 0.76 cm, with the edge width changed correspondingly to keep the same porosity, were fabricated to investigate the effect of cell size on compressive properties and build time. The yield strengths and compressive moduli obtained from the compression tests are shown in Fig. 13 and Fig. 14, respectively.

The yield strengths and compressive moduli obtained from the compression tests are shown in Fig. 13 and Fig. 14, respectively. The data indicate that yield strength decreases with increase in edge length between 0.31 cm and 0.61 cm, but then it increases when the edge length increases from 0.61 cm to 0.76 cm. The highest yield strength occurs at 0.76 cm edge length. The compressive modulus obtained from the honeycomb specimens with the various edge lengths are shown in Fig. 14. The compressive modulus decreases with increase in edge length when the edge length increases from 0.31 cm to 0.76 cm.) a P M( h t g n e r t Sd l e i Y38 37 36 35 34 33 32 31 3036.0435.2934.6836.710.20.30.40.50.60.70.80.9Edge Length, (cm)Figure 13. Comparison of yield strength vs.

Comparison of yield strength vs. edge length for honeycomb strucutres with the same porosity.a P M1,7001,5001,521.7, s u l u d o M e v i s s e r p m o C1,3001,1009007005001,217.2914.7610.20.240.320.400.480.560.640.720.80Edge Length, cmFigure 14. Compressive modulus vs. edge length for honeycomb structures with the same porosityThe stress-strain curves for the honeycomb specimens with varying edge lengths (with the corresponding changes in edge width to keep the same porosity) are shown in Fig. 15. All of the stress-stain curves have similar trends, and the 0.76 cm edge length has the maximum strength at approximately 3.8% strain. Figure 16 provides the stress-strain curves with more details in the range of 3% to 6% strain for the various edge lengths.a P M, s s e r t S45 40 35 30 25 20 15 10 5 00.31 cm 0.46 cm 0.61 cm 0.76 cm0369 12 15 18 21 24 27 30Strain, %Figure 15.

Stress-strain curves for various honeycomb structures with different edge lengths at the same porosity4443420.31 cm 0.46 cm 0.61 cm 0.76 cma P M41, s s e r t S40393837363456Strain, %Figure 16. Stress-strain curves in the range of 3% to 6% strain for various honeycomb specimens with different edge lengths at the same porosityThe build times for the various honeycomb specimens that vary in edge length at the same porosity are shown in Fig. 17, which shows that build time decreases with increase in edge length. The build time of the specimen having 0.76 cm in edge length is 36.5% shorter than the specimen having 0.31 cm in edge length. By combining the data in Figs.

The build time of the specimen having 0.76 cm in edge length is 36.5% shorter than the specimen having 0.31 cm in edge length. By combining the data in Figs. 15 and 17, we see that among the four different cell sizes (0.31, 0.46, 0.61, and 0.76 cm in edge length) of the honeycomb structure with 57% in porosity, the cell size with 0.76 cm edge length has not only the highest yield strength but also the shortest build time.70n i m e m T d l i u B,i605040302063.064.055.040.01000.20.40.60.8Edge Length, cmFigure 17. Build time vs.

Build time vs. edge length for the honeycomb structureConclusionsExperimental studies were conducted to evaluate and compare the compressive properties and build time for periodic lattice structures that have unit cells including honeycomb, square, diamond, circle, and triangle shapes, as well as the sparse and sparse-double dense build styles, fabricated by the fused deposition modeling (FDM) process using Stratasys Fortus 400 machine. The results are summarized below:1) The honeycomb structure has higher yield strength and compressive modulus than the other four lattice structures, and all of the five lattice structures evaluated have much higher yield strength and compressive modulus than the sparse and sparse-double dense build styles. However, the build times of the five lattice structures are approximately twice as much as the build times of the sparse and sparse-double dense builds.

However, the build times of the five lattice structures are approximately twice as much as the build times of the sparse and sparse-double dense builds. Thus, there is a clear trade-off between the cellular lattice structures and the sparse and sparse- double dense builds in terms of compressive properties and build time.2) The compressive properties and build time depend on part porosity for a given lattice structure. The yield strength and compressive modulus decrease and the build time decreases when the part porosity with the honeycomb structure increases. When the porosity of honeycomb specimen increases from 57% to 71% in porosity (~25% increase), correspondingly there is an increase in yield strength by 57%, and a decrease in build time by 46%.3) The compressive properties and build time also depend on the cell size for a given lattice structure with a constant porosity.

For the honeycomb structure having the edge length ranging from 0.31 cm to 0.76 cm with the edge width changed correspondingly to keep the same porosity, the yield strength decreases with increase in edge length between 0.31 cm and 0.61 cm but then increases with increase in edge length from 0.61 cm to 0.76 cm. The highest yield strength occurs at 0.76 cm edge length. The compressive modulus decreases continuously when the edge length increases from 0.31 cm to 0.76 cm.

re explored to overcome this problem [6]. Selective Laser Sintering was applied in ceramic part fabrication by mixing organic binder (typically PMMA) with ceramic particles as the process material [7], but the post-sintering density is relatively low at about 55% [8].Most SFF techniques for ceramic component fabrication involve the use of organicbinders. In some processes, such as FDC, the binder content may be as high as 40 to 50 vol.%. This organic binder needs to be removed during post processing and generates harmful wastes that are undesirable for the environment [9]. Freeze-form Extrusion Fabrication (FEF) extended the idea of the rapid freeze prototyping (RFP) method [10-13], where water droplets are deposited on demand and freeze on a 2-D motion substrate for the fabrication of 3D components in a layer-by-layer manner. FEF uses an aqueous ceramic paste with a solids loading up to 50 vol.%. The organic binder content is only 2-4 vol. %.

FEF uses an aqueous ceramic paste with a solids loading up to 50 vol.%. The organic binder content is only 2-4 vol. %. In FEF, aqueous-based ceramic paste is extruded using a ram extruder and deposited on a 2-D motion substrate. After the deposition of each layer the Z-axis moves up by one layer thickness and the next layer is deposited. This process is repeated until the component is completely fabricated. Freeze-drying is used to prevent crack formation during the drying process. After freeze-drying, the binder is removed in a rapid heating cycle because of the low binder content. Finally, the parts are sintered at 1550 °C. The post processing has been detailed previously [9]. FEF has some unique advantages, such as large component fabrication, high density of sintered components, low percentage of organic binder, and almost no material waste.

FEF has some unique advantages, such as large component fabrication, high density of sintered components, low percentage of organic binder, and almost no material waste. FEF is an environmentally friendly SFF technique.In the present paper, a custom-deigned 3D gantry system with a custom-designed coolingsystem (0 °C to -30 °C) was used for FEF processing. The layer thickness optimization was performed for fabrication at -20 °C. The self-sustaining ability of the process at different temperatures was analyzed. The time constant and gain of the first-order process model of FEF were calculated. The trends of these two parameters during fabrication at -20 °C and at 0 °C were recorded and analyzed.2. Experimental Setup and Procedure2.1 Experimental setupThe 3-D gantry system, as shown in Figure 1, consists of three orthogonal linear axesfrom Velmex BiSlide (Velmex, Bloomfield, NY), each with a 508 mm travel range.

The X-axis consists of two parallel slides and is used as the support for the Y-axis. The two parallel slides provide smooth and stable motion and allow more fabrication space. The Z-axis is mounted on the Y-axis and the extrusion mechanism is mounted on the Z-axis. All axes are mounted with limit switches on both ends. Four DC motors (Pacific Scientific PMA22B), each with a resolver for position feedback at a resolution of 1000 counts per revolution drive the axes. Each motion axis has a maximum speed of 127 mm/s and a resolution of 0.00254 mm. All the axes are controlled by a Delta-Tau Turbo PMAC PCI board.The right image in Figure 1 shows an enlarged view of the extrusion mechanism. It is aram extruder driven by a DC motor (Kollmorgen AKM23D), which has an encoder with a resolution of 0.254 µm and is mounted on the Z-axis slide. An Omega LC305-1KA load cell (Danaher Motion, Wood Dale, IL) is mounted between the plunger and the ram extruder tomeasure the extrusion force.

An Omega LC305-1KA load cell (Danaher Motion, Wood Dale, IL) is mounted between the plunger and the ram extruder tomeasure the extrusion force. An analog-to-digital conversion board (Delta-Tau ACC28) converts the analog signal from the load cell to digital signal and inputs to the PMAC board.The 3-D gantry is located inside a freezer. A condenser is used to keep the freezer temperature at 0 °C (± 2 °C). Liquid nitrogen is used for lowering the temperature of the freezer between 0 °C to –30 °C. A temperature controller (Omega CN132) (Danaher Motion, Wood Dale, IL) is used to control the temperature of the freezer through a solenoid valve, which regulates the flow rate of nitrogen.

A temperature controller (Omega CN132) (Danaher Motion, Wood Dale, IL) is used to control the temperature of the freezer through a solenoid valve, which regulates the flow rate of nitrogen. As shown in Figure 2, two heating coils are installed around the extruder and the nozzle to keep the paste temperature in the range of 10-15 °C to prevent the paste from freezing.Z axisRam extruderY axisLoad cellX-X axisFigure 1: The 3D gantry system with extrusion mechanism.Syringe holderHeating coilsSyringeHeat insulationHeating coil DC power supplyFigure 2: Schematic drawing showing the syringe and nozzle heating system.2.2 Process parametersThe process parameters include initial extrusion force, extrusion force increment, roadoffset, layer thickness, and X-Y motion table (working surface) speed. The road offset between deposition trajectories was mainly determined by the diameter of the nozzle. A 580 µm diameter nozzle was used in all experiments.

The road offset between deposition trajectories was mainly determined by the diameter of the nozzle. A 580 µm diameter nozzle was used in all experiments. The X-Y motion speed was in a range of 10-20 mm/s. The X-Y table speed was set at a certain speed and the layer thickness was adjusted according to nozzle size and fabrication temperature.A larger extrusion force is associated with a higher extrusion rate. The initial extrusionforce was set to 311 N. To maintain a constant extrusion rate and avoid nozzle clogging, the reference extrusion force was continually increased slowly.

The initial extrusionforce was set to 311 N. To maintain a constant extrusion rate and avoid nozzle clogging, the reference extrusion force was continually increased slowly. An adaptive PI controller was designed and implemented to control the ram velocity in a range of ± 50 µm/s at steady state to achieve the desired extrusion.2.3 Investigation of fabrication at different temperatures2.3.1 Layer thickness effectThe layer thickness determined by the Z-axis shift distance needs to be carefully adjusted,so that the nozzle does not disturb the previous layer during the deposition process. To optimize this parameter, cylinder samples were fabricated at 0 °C and -20 °C with different layer thicknesses and X-Y table speeds as listed in Table 1.

To optimize this parameter, cylinder samples were fabricated at 0 °C and -20 °C with different layer thicknesses and X-Y table speeds as listed in Table 1. The table speed for fabrication at -20 °C was slightly larger than at 0 °C to compensate for the temperature effect of the paste.Table 1: Deposition parameters used in the layer thickness experiments1Temperature (ºC) 0X-Y table speed (mm/s) 10Layer thickness (µm) 5102201151032011580420116402.3.2 Minimum deposit angle testThe minimum deposit angle is the minimum angle that can be achieved between thesubstrate and the slope of a hollow cone without collapse. This angle reflects the offset ability of the FEF process in building a 3-D part without use of support material. Three sets of tests were conducted to fabricate cones with different bottom diameters to find the minimum deposit angle. The tested temperatures include 20 °C, 10 °C, 0 °C, -10 °C, and -20 °C.

Three sets of tests were conducted to fabricate cones with different bottom diameters to find the minimum deposit angle. The tested temperatures include 20 °C, 10 °C, 0 °C, -10 °C, and -20 °C. In each set of tests, hollow cones were fabricated using bottom diameters of 38 mm, 51 mm, and 64 mm. The cone height was varied to determine the lowest height without collapse of the cone during the building process for calculation of the minimum deposit angle.2.3.3 Time constant and gainThe FEF process contains many nonlinear effects, such as air bubbles trapped within the paste, uneven water content in different portions of the paste, various sizes of agglomerates, etc. Paste consistency is also slightly different from batch to batch. These factors contribute to difficulties in modeling the paste extrusion process.

Paste consistency is also slightly different from batch to batch. These factors contribute to difficulties in modeling the paste extrusion process. Previous research work shows that theextrusion process, in general, can be approximated as a first-order dynamic system, where commanded voltage to the ram motor amplifier is the input and extrusion force is the output [10]. However, according to the experimental data, there is significant variation in the model parameters during the extrusion. Therefore, the Recursive Least Square (RLS) method is applied to model the extrusion process and to determine how the amount of remaining paste in the material reservoir affects the dynamic model parameters. The model input is the commanded ram motor voltage, which is processed by a 16 bit digital-to-analog converter before going into the ram motor amplifier. The output is the measured extrusion force, which is sent to the computer via an analog-to-digital converter.

The output is the measured extrusion force, which is sent to the computer via an analog-to-digital converter. The resolution of the measured extrusion force is 2.224 N.Experiments were conducted to investigate how the parameters of the extrusion dynamicmodel varied during extrusion. In these experiments, a command ram motor voltage of 3 mV is sent to the motor amplifier for 10 seconds, and then the voltage is changed to -1.5 mV for 5 seconds, increased to 3.75 mV for another 10 seconds, and decreased to -1.75 mV for 5 seconds. Each test lasted until approximately 5 ml of paste was extruded. The extrusion force was measured with a sample period of 0.06 sec, and the RLS method was applied to estimate the model parameters at each sample period. The experiments were repeated at two environment temperatures.3.

The experiments were repeated at two environment temperatures.3. Results and Discussion3.1 Relationship between extrusion force and extrusion rateDuring fabrication the extrusion rate decreased as the amount of paste in the syringe decreased when the extrusion force was constant. The table speed was maintained constant during fabrication. Therefore, the decrease in extrusion rate may result in under-filling and generating scraps in the building area as shown in Figure 3. Although the reason for this phenomenon is still not entirely clear, it is believed that this phenomenon is related to the paste property changing during the extrusion. There is a redistribution of liquid phase versus solid phase in the paste as the extrusion process proceeds.The extrusion force should be gradually increased to balance the gradual increase of thepaste resistance force during extrusion. The optimal rate of extrusion force increase during fabrication was approximately 2.3×10-2 N/s.

The optimal rate of extrusion force increase during fabrication was approximately 2.3×10-2 N/s. By using an adaptive PI controller at a sample period of 0.06 seconds, the extrusion force followed the continually increasing reference force closely as shown in Figure 4.ScrapsFigure 3: Picture showing scraps on a tangent ogive hollow cone.440420Actual force Reference force400)N(380e c r o f360m a R3403203002800100020003000 Time (second)400050006000Figure 4: Reference force vs. actual ram force during an extrusion process.3.2 Temperature effectsAt 20 °C environment temperature, the extruded material drying rate was found to befairly high if the moisture was less than 60%. The high drying rate was usually associated with non-uniform drying, which generated cracks and caused warping in the fabricated part.

The high drying rate was usually associated with non-uniform drying, which generated cracks and caused warping in the fabricated part. Increasing moisture in the surrounding area could help in improving sample quality.At 0 °C, water evaporated more slowly and the drying rate is much lower than that at 20 °C. The surface finish of the samples fabricated at this temperature was generally smoother than the surface finish of those fabricated at 20 °C.Because the extruded materials dried slowly and didn’t freeze, the lower portion of thecomponent did not have enough strength to support the whole component toward the end of the fabrication for a large component. This may lead to component deformation or even collapse. Decreasing the table speed (< 12.5 mm/s), which increases drying time, could help solve this problem.

This may lead to component deformation or even collapse. Decreasing the table speed (< 12.5 mm/s), which increases drying time, could help solve this problem. However, the slow deposition rate would increase the fabrication time, which is undesirable.At -20 °C, heating coils (Figure 2) were used to keep the paste warm at 10-15 °C to ensure proper extrusion. The extruded ceramic paste could freeze at -20 °C. No visible part deformation was observed. The table speed used for fabrication at -20 °C was 15-20 mm/s, while 10-12 mm/s was used for fabrication at 0 °C.

No visible part deformation was observed. The table speed used for fabrication at -20 °C was 15-20 mm/s, while 10-12 mm/s was used for fabrication at 0 °C. Table 2 shows a general comparison of fabrication at these three temperatures.Table 2: Results of layer thickness experimentsEnvironment Temperature (ºC)Dry RateFreeze RateTable Speed (mm/s)Part DeformationPart Surface20LowHigh15-20NoSmooth0MediumNone10-12YesSmooth20High, non-uniformNone10-12YesCracks and warping3.3 Comparison of part fabrication at different temperatures3.3.1 Layer thickness effectFive cylinders were fabricated for each layer thickness. Figure 5 shows one typicalcylinder for each value listed in Table 1. Cylinder 1 was fabricated at 0 °C, while cylinders 2-4 were fabricated at -20 °C. Visually cylinder 1 has the best surface finish. This is because at this temperature, the extruded paste had a high viscosity (>50 Pa.s).

Visually cylinder 1 has the best surface finish. This is because at this temperature, the extruded paste had a high viscosity (>50 Pa.s). As a result, the new layer of material deforms slightly, making the surface smoother. The layer thickness for this fabrication was 510 µm, which was slightly less than the diameter of the nozzle (580 µm). At –20 °C, the paste froze quickly and exhibited little deformation. When using the same layer thickness (510 µm), the surface quality was poor as shown in Figure 5 (cylinder 2). This is because the paste material did not deform and the nozzle tip scratched the previously deposited materials. By increasing the layer thickness to 580 µm, the cylinder (Figure 5, cylinder 3) had a better surface finish. When the layer thickness was increased to 640 µm, the cylinder (Figure 5, cylinder 4) had an even better surface finish. However, when the layer thickness was larger than 640 µm, under- filling occurred.

However, when the layer thickness was larger than 640 µm, under- filling occurred. Therefore, the optimal layer thickness should be 640 µm for fabrication at -20 °C.1 (0°C, 510 µm) 2 (-20°C, 510 µm) 3 (-20°C, 580 µm) 4 (-20°C, 640 µm)Figure 5: Cylinders fabricated using the parameters in Table 13.3.2 Minimum deposit angle testFigure 6 gives the definition of the minimum deposition angle. Figure 7 shows a successfully built cone with 19 mm height and a 38 mm bottom diameter (left) and a collapsed cone (right).The minimum deposit angle test results are given in Table 3. The relationships between the minimum deposit angle and the fabrication temperature are plotted in Figure 8. The figure shows that the minimum deposit angle increases quickly from -20 °C to 0 °C, then more slowly from 0 °C to 10 °C, and the trend further flattens from 10 °C to 20 °C. In the -20 °C experiment, the extruded materials froze and became solid quickly.

In the -20 °C experiment, the extruded materials froze and became solid quickly. This is the reason for the smallest minimum deposit angle at this temperature for the range between -20 °C and 20 °C. When the temperature increased to 0 °C, the extruded materials would not freeze but the viscosity was high.

The model parameters (i.e., time constant τ and gain K) are derived from the estimated parameters a and b and, respectively, areτ = −T lnaK=1b a −(1)(2)(4)(5)(6)(7)3.4 Demonstration of components fabricationFigure 10 shows tangent ogive cones in green state fabricated by the FEF process at -20°C environment temperature. The left image shows a cone made from alumina paste and the right image shows two cones made from zirconium diboride.25mm25mmFigure 10: One Al2O3 and two ZrB2 tangent ogive hollow cones in green state4. Summary and ConclusionsThe environment temperature has been found to significantly affect the material extrusionand deformation behavior in the freeze-from extrusion fabrication (FEF) process using aqueous- based ceramic pastes. Fabricating samples at -20 °C allows the deposited paste to freeze during the fabrication process.

Fabricating samples at -20 °C allows the deposited paste to freeze during the fabrication process. This results in the lowest minimum deposit angle, which means the highest self-sustaining ability (without the use of support material). The minimum deposit angle is strongly related to the environment temperature. The lower the environmental temperature is, the smaller the minimum deposit angle will be. Thus the FEF process can fabricate larger parts at -20 °C than at 0 °C. By using a heater to prevent the paste from freezing, the model time constant and the model gain are both smaller at -20 °C than the values of these parameters at 0 °C. Therefore, at -20 °C, the paste with heating is easier to extrude, and the system response is faster than at 0 °C without heating the paste.

Therefore, at -20 °C, the paste with heating is easier to extrude, and the system response is faster than at 0 °C without heating the paste. The successful fabrication of tangent ogive cones at -20 °C from alumina and zirconium diboride pastes have demonstrated the feasibility of the FEF process.AcknowledgementThis work was supported by the Air Force Research Laboratory under Contract FA8650-04-C-5704.References1. T. S. Huang, “Fabrication of Ceramic Components Using Freeze-form Extrusion Fabrication,” Ph.D. Dissertation, Department of Materials Science and Engineering, University of Missouri-Rolla (2007).2. G. M. Lous, I. A. Cornejo, T. F. McNlty, A. Safari, and S. C. Danforth, “Fabrication of Piezoelectric Ceramic/Polymer Composite Transducers Using Fused Deposition of Ceramics,” J. Am. Ceram. Soc., 83 [1] 124-28 (2000).3. J. J. McIntosh, S. C. Danforth and V. R.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEXPERIMENTAL INVESTIGATION OF FLUID-PARTICLE INTERACTION IN BINDER JET 3D PRINTINGJoshua J. Wagner*, Hang Shu*, Rahul Kilambi*, and C. Fred Higgs III*Department of Mechanical Engineering, Rice University, Houston, TX 77005ABSTRACTWide-scale adoption of binder jet 3D printing for mission-critical components in aerospace, biomedical, defense, and energy applications requires improvement in mechanical properties and performance characteristics of end-use components. Increased fidelity may be achieved with better understanding of the interfacial physics and complex fluid-particle interactions fundamental to the process. In this work, an experimental testing apparatus and procedure is developed to investigate the fluid and particle dynamics occurring upon impact of jetted binder droplets onto a powder bed.

In this work, an experimental testing apparatus and procedure is developed to investigate the fluid and particle dynamics occurring upon impact of jetted binder droplets onto a powder bed. High-speed, microscopic imaging is employed to capture short time-scale phenomena such as ballistic particle ejection, capillary flow, and particle clustering. The effects of different process parameters (e.g., translational printhead velocity, jetting frequency, and impact velocity) on the dynamics of Inconel powder are studied. These experiments reveal that the fluid-particle interaction is significantly affected by a combination of printing parameters, ultimately governing the quality and performance of binder jet 3D printed components.1. INTRODUCTIONBinder jet 3D printing (BJ3DP) is a type of powder bed additive manufacturing (AM) process that utilizes a scanning printhead to selectively administer fluid binder into a powder layer to form a cross-section of a three-dimensional object.

After completion, the components are buried in unbound powder (Fig. 1e); therefore, a depowdering step is required for retrieval. Typically, the printed parts are then either sintered or infiltrated with a secondary material in a high-temperature furnace for densification and strengthening [2]. An extensive review of BJ3DP, from its development to present state-of-the-art, is given in [3].The underlying physics of the BJ3DP process allows for many unique advantages compared to other AM methods, such as a wide range of suitable materials, relatively fast build times, and cost-effectiveness. These advantages have generated interest in BJ3DP for many applications in a variety of industrial sectors; however, its end-use components generally have134(a)(b)(c)(d)(e)FIGURE 1. The general BJ3DP process comprised of the following stages: (a) powder delivery, (b) powder spreading, (c) binder jetting, and (d) build platform lowering.

The general BJ3DP process comprised of the following stages: (a) powder delivery, (b) powder spreading, (c) binder jetting, and (d) build platform lowering. The stages (a)-(b) are repeated until the complete geometries are formed (e). Note that hatched regions indicate unbound powder.less robust mechanical increased porosity and properties compared to powder bed fusion type AM processes. These downfalls have limited the use of BJ3DP for the fabrication of high-performance parts in demanding applications [4], e.g., combustion chambers for liquid rocket engines [5] or pressure vessel components in nuclear fusion reactors [6]. is of vital Therefore, mitigating issues importance facilitate a more wide-spread implementation of BJ3DP among a range of industries.thesetoTo achieve this, many research efforts have been carried out in each major stage of the process. Several powder dispensing (Fig.

Several powder dispensing (Fig. 1a) techniques have been explored for use in powder bed AM methods – a review of which is given in [7]. Powder spreading (Fig. 1b) simulations were performed using discrete element modeling (DEM) to better understand how spreading parameters affect the quality of the layer [8, 9]. The binder jetting stage (Fig. 1c) has been studied by a number of researchers; several of the important works on this topic are presented in the following section. Finally, there have been efforts to improve the various post-processing techniques often used to strengthen binder jetted parts, e.g., see [10, 11, 12].The goal of this work is to develop an experimental apparatus and testing procedure that allows for investigation of the interfacial fluid- particle interaction prevalent during the binder jetting stage.

Ultimately, this platform will provide insight on how printing parameters can be tuned to achieve optimal printing characteristics for a particular feedstock powder. The apparatus is constructed to replicate the fundamental printing operations in such a way that allows for quantative measurements of the print quality depending on various input parameter combinations. High-speed, microscopic imaging is employed to capture short time-scale dynamics occurring upon the impact of jetted binder droplets onto a powder bed. A dimensional analysis of the process reveals pertinent dimensionless groups that influence the coupled fluid and particle dynamics. Experiments are carried out to determine the effects135that certain dimensionless groups have on powder binding behavior.2.

Experiments are carried out to determine the effects135that certain dimensionless groups have on powder binding behavior.2. FLUID-PARTICLE INTERACTION IN BJ3DPThe instant a jetted binder droplet impinges the powder bed, many physical phenomena begin to take place that govern the formation of bound particle features that make up the part. Immediately upon impact, momentum transferred from the droplet to the unconsolidated powder bed can result in ballistic ejection of particles. This potentially gives rise to various types of part defects; for example, the ejection of particles yield impact craters which are deleterious to dimensional accuracy and serve to increase porosity [13].At impact, inertial forces initially drive the flow of liquid either into or on top of the bed. As the initial kinetic energy of the drop is transferred and dissipated into the powder bed, the binder may continue to flow into the interstitial space between particles due to capillary forces [14].

Simultaneously, the liquid attempts to minimize its surface energy by generating interfacial forces which impede the expansion of surface area as the binder spreads. These interfacial effects, namely surface tension and capillary action, often produce a net attractive force between mutually wetted particles. If these forces are large enough to overcome particle inertia and friction, the particles can be accelerated and drawn inward towards each other, forming a bound cluster referred to as a “primitive” [15].A single-drop primitive is the most basic building element produced in BJ3DP. The printhead translates over the powder bed at a given velocity where, ideally, successive droplets coalesce with the previous ones forming a single-line primitive [16]. The line primitives must then be adequately stitched to adjacent lines to form a solid cross-section of the part. Additionally, the liquid should penetrate the powder bed deep enough to achieve sufficient inter-layer binding.

Additionally, the liquid should penetrate the powder bed deep enough to achieve sufficient inter-layer binding. For these reasons, the formation of line primitives is a predominant factor governing the relative density, dimensional accuracy, resolution, surface properties, and overall quality of the parts.In Ref. [29], high-speed imaging of the fluid-particle interaction in BJ3DP was performed with synchrotron X-rays using the Advanced Photon Source at Argonne National Laboratory. With this approach, an X-ray beam passes completely through the powder bed and is converted to visible light on the opposite side. This approach has provided new insights into the surface and sub-surface dynamics of the powder bed during binder jetting. At the powder bed surface, ballistic ejection, impact cratering, and primitive formation were observed for several powder types of varying particle size and morphology.

At the powder bed surface, ballistic ejection, impact cratering, and primitive formation were observed for several powder types of varying particle size and morphology. Powder bed compression induced by the droplet impact was discovered, and the depth of disturbance for each powder was reported. Although this work has offered knowledge about the effect of particle size and shape on the binder-powder interactions, the influence of printing parameters on primitive formation remains poorly understood.

(1) expresses the ratio of binder volume to void volume in a predefined envelope (cid:4666)(cid:543)(cid:1876) (cid:3400) (cid:543)(cid:1877) (cid:3400) (cid:543)(cid:1878)(cid:4667) within the powder bed. Saturation is an important parameter that must be optimized for particular feedstocks. Typically, higher saturation results in increased green (pre-sintered) part strength; however, excessive binder will bleed out of the desired part boundaries reducing dimensional accuracy and resolution [14]. Presently, the optimal saturation for a given powder is most often found by trial and error; however, several efforts have been made to obtain it more rigorously. In Ref. [18], an analytical model was developed to estimate the saturation level that results in a balance of capillary pressure dependent on contact angle of the solid-liquid-gas interface, binder-air surface tension, and powder bed packing fraction.

This “equilibrium saturation” was first suggested to be the optimal saturation for BJ3DP in Ref. [14] based on arguments from the classical theory of fluid flow in porous media [19]. A more direct experimental approach was taken in Ref. [20] by analyzing the influence of saturation on the dimensional accuracy and surface roughness of 420 stainless steel parts. The same methodology was adopted in Ref. [21] to investigate how saturation affects the flexural strength of 316L stainless steel parts. Further studies regarding the appropriate selection of saturation level for various powders were carried out in Refs. [22, 23, 24, 25].While some success in improving part quality was achieved in the above works, saturation alone is not sufficient to fully characterize the effects of fluid-particle interaction in BJ3DP. This is primarily due to the fact that saturation is not a unique value, which is made clear by closer inspection of Eq. (1).

This is primarily due to the fact that saturation is not a unique value, which is made clear by closer inspection of Eq. (1). An infinite combination of the variables on the right-hand side may produce the same saturation value; however, holding saturation constant while changing other parameters does not guarantee that primitive formation behavior will remain the same. For example, a (cid:3) , can be accompanied with a thinner layer thickness, (cid:543)(cid:1878), reduction of binder droplet diameter, (cid:1830)(cid:3029) (cid:3) will also reduce the kinetic energy of the jetted to produce the same value of (cid:1845), but a decreased (cid:1830)(cid:3029) droplet, which can dramatically change the impact dynamics. In this work, a dimensional analysis is performed to reveal how important parameters are related and how they should be varied in order to carry out valid parametric studies (Sec. 3).3.

In this work, a dimensional analysis is performed to reveal how important parameters are related and how they should be varied in order to carry out valid parametric studies (Sec. 3).3. EXPERIMENTAL SETUPA testing apparatus was constructed to allow for close investigation of fluid-particle interaction in BJ3DP and the effect printing parameters have on the binding of a single-line primitive. The apparatus (Fig. 2) is comprised of a linear motor that drives a single-jet droplet dispenser over a powder sample at an adjustable translation speed, (cid:1847)(cid:3031), along an x-axis rail. The x- axis motor and rail assembly is mounted on a linear stage that allows for precise manual adjustment of the dispenser’s y-axis position.

(b) Magnified diagram of the single-jet droplet dispenser and powder sample.frequency, (cid:1858)(cid:3037), and droplet speed,(cid:3)(cid:1847)(cid:3037)(cid:481) along the jet trajectory. The 53 μm orifice on the dispenser nozzle produces droplet sizes on the order of 50 μm. The droplet diameter can be slightly adjusted by varying the parameters of piezoelectric actuation.The powder sample is contained in a cylindrical vessel having a diameter and height of 1 cm. In this work, the fundamental phenomenon of line primitive formation on the surface of an unbound powder bed is considered; therefore, a thin layer thickness is not currently imposed. After the powder is loaded, the vessel is lightly tapped and the powder bed surface is smoothed by manually spreading the excess powder with a lab spatula. Future work will examine the effect of layer thickness on the line primitive and will replicate the BJ3DP spreading process to greater fidelity.

Future work will examine the effect of layer thickness on the line primitive and will replicate the BJ3DP spreading process to greater fidelity. Highly spherical, plasma atomized Inconel 718 powder is used. The particle diameter range is 20-38 μm and the tapped packing fraction (cid:2038) is approximately 0.53. Inconel is a nickel- chromium-based superalloy that maintains high strength at extreme temperatures and has excellent oxidation and corrosion resistance [26]. For these reasons, Inconel alloy components are commonly used in gas-turbine engines, combustion chambers, nuclear reactors, and other demanding applications involving high thermal loading. Currently, BJ3DP has been employed to manufacture components from Inconel 625 and Inconel 718 powder feedstocks [27, 28]; however,138further work is required to mitigate part defects introduced by the process before the components can be used in mission- critical and high-performance applications..

(cid:2010) ' ' ' ' \' ' ' ' 'A Vision Research Phantom VEO high-speed camera with a 4.5X microscope objective is mounted above the powder sample and droplet dispenser assembly. Various camera angles were attempted such that the line of sight of the camera intersected the powder bed at an oblique angle rather than perpendicular to it; however, because of the shallow depth of field at high magnification, a purely top-down arrangement with the camera’s line of sight parallel to the z- axis (as it is in Fig. 2a) produced the best results. In this configuration, view of the droplet impact location on the powder bed is obstructed by the single-jet dispenser (Fig. 3a). This issue is addressed by positioning the dispenser with a slight angle (cid:2010) about the y-axis in order to move the droplet impact location in the line of sight of the camera (Fig. 3b). Assuming the translational dispenser motion is from right to left in Fig.

The basic operation of synchronized stroboscopic illumination is shown in Fig. 4. The strobe is illuminated at a frequency that is synchronized with the piezoelectric actuation of the dispenser. A small time delay (cid:2028) is imposed between the jet initiation and LED pulses; thus, a low framerate camera (typically 24-60 frames per second) will only capture what appears to be a single stationary droplet at some position along its trajectory [30]. The distance of the “pseudo-stationary” droplet from the dispenser orifice, (cid:1856)(cid:3043)(cid:3046), and the time delay can be used to calculate the jetting velocity as (cid:1847)(cid:3037) (cid:3404) (cid:1856)(cid:3043)(cid:3046)(cid:512)(cid:2028). Furthermore, the droplet diameter is also easily obtained the stroboscopic images.

Due to the small volume of binder jetted (~50 picoliters per drop), it can be difficult to discern the liquid within the bed, even when high magnification is used. This issue is made less severe by introducing a small amount of fluorescent dye to the binder prior to jetting. This causes the liquid contained within the line primitives to illuminate upon exposure to ultraviolet light.4. DIMENSIONAL ANALYSISA dimensional analysis is carried out to aid in the design of experiments of single-line primitive formation. Dimensional analysis is a powerful technique that can reduce the number of independent variables in the problem by grouping them into a fundamental set of nondimensional parameters for which the problem depends [31]. To do this, all parameters that can affect the fluid- particle interaction in a single-line print must be considered. The problem being analyzed is that of a continuous droplet train impacting a powder bed.

The problem being analyzed is that of a continuous droplet train impacting a powder bed. The droplets are generated from a single-jet dispensing device translating over a powder bed at a constant speed. Each droplet is assumed to be spherical and of fixed diameter – satellite droplets are not currently considered. Additionally, temperature is assumed to only affect the material properties of the binder and powder. Although their influence may be significant, evaporation and other phase change effects are not considered here. Finally, the dispenser angle (cid:2010) is set to zero; therefore, (cid:1873)(cid:3051) (cid:3404) (cid:1847)(cid:3031) and (cid:1873)(cid:3053) (cid:3404) (cid:1847)(cid:3037). Given the problem statement and the aforementioned assumptions, the relevant parameters and their nomenclature are given in Table 1.TABLE 1. Nomenclature and important parameters in a single-line primitive formation event.

two this response variables are considered: 1) the fractional area coverage, (cid:1827)(cid:1832), of a line primitive contour with its bounding box, and 2) the bounding box width (cid:1875). As shown in Fig. 5, the red bounding box is constructed from the four extreme points (top, bottom, left, and right) of the 2D primitive line contour. The area fraction is calculated by (cid:1827)(cid:1832) (cid:3404) (cid:1827)(cid:3039)(cid:3043)(cid:512)(cid:4666)(cid:1864) (cid:942) (cid:1875)(cid:4667), where (cid:1827)(cid:3039)(cid:3043) is the gray region representing the bound line primitive in Fig. 5. It is clear from Eq. (5) that (cid:1844) is some function of the fourteen independent parameters given in Table 1. With Buckingham’s Pi theorem, Eq. (5) is cast into another functional relationship between the nondimensional groups that can be constructed from all of the independent and dependent variables:determinationl(cid:1875)Instudy,(cid:1864)FIGURE 5.

This issue is presently being addressed and is deferred to later studies.Still frames extracted from the high-speed videos are shown in Fig. 6. The x-component of the droplet impact velocity for this case is (cid:1873)(cid:3051) (cid:3404) (cid:882)(cid:484)(cid:883)(cid:884)(cid:889) m/s. In agreeance with the findings in [29], a significant number of ballistic particle ejections were observed. Furthermore, several tests were performed with a lower vertical impact speed ((cid:1873)(cid:3053) (cid:3404) (cid:884)(cid:484)(cid:891)(cid:887) m/s) and particle ejections were still prevalent, albeit with reduced occurrences. For lower vertical impact speeds, the line primitives were found to be more susceptible to breakup. This is attributed to the “balling” phenomenon, which is the result of the Plateau-Rayleigh instability in liquid filaments [32]. Balling in BJ3DP was discussed in Refs.

This is attributed to the “balling” phenomenon, which is the result of the Plateau-Rayleigh instability in liquid filaments [32]. Balling in BJ3DP was discussed in Refs. [3, 17].Four trials are performed utilizing the still frame imaging setup for a (cid:1873)(cid:3051) range of (cid:882)(cid:484)(cid:882)(cid:883)(cid:888)-0.127 m/s. The resulting line primitives are shown in Figs. 7-10 in order of decreasing (cid:555)(cid:2869)(cid:2869) ((cid:1845)(cid:1872)) and (cid:555)(cid:2869)(cid:2870)(cid:3)(cid:4666)(cid:1873)(cid:3051)(cid:512) (cid:1873)(cid:3053)(cid:4667). The line primitive contours are generated using image processing software developed in-house. Due to the illumination of the fluorescent binder, the line primitive edges are clearly line primitive shape and bounding box for each trial is given in Fig. 11. Unsurprisingly, from Figs. 7-11, it is evident that the bounding box width increases as (cid:555)(cid:2869)(cid:2869) is decreased.

11. Unsurprisingly, from Figs. 7-11, it is evident that the bounding box width increases as (cid:555)(cid:2869)(cid:2869) is decreased. This is attributed to the fact that (cid:555)(cid:2869)(cid:2869) is the the ratio of droplet inverse of spacing (Eq. 2) to droplet diameter; thus, a lower value indicates a larger volume of binder being deposited during the dispenser translation. In addition to the width increase, (cid:1827)(cid:1832)(cid:3)also increases with decreasing (cid:555)(cid:2869)(cid:2869). The results for the four trials are given in Table 3.delineated.TheFIGURE 6. Still frames captured from high-speed video of the droplet dispenser traversing from left to right. The frames are in downward sequential order.142FIGURE 7.

The characteristics of the line primitives that make up the part largely influence its mechanical properties, dimensional accuracy, resolution, and overall quality. This work will aid in the improvement of such attributes by providing an experimental platform to investigate the relationship between the binding process parameters and the resulting144-,---line primitives. With this insight, an optimal combination of process parameters that yield the highest quality components may be discovered for any powder feedstock.In this study, the apparatus was employed to investigate single-line primitive formation on Inconel powder at various translational dispenser speeds. The area fraction and width of the line primitives were both shown to increase with decreasing translational speed. In future work, the test platform described here will be employed to perform extensive parametric studies on line primitive formation for various powder samples.

In future work, the test platform described here will be employed to perform extensive parametric studies on line primitive formation for various powder samples. The final parts will then be examined to reveal whether superior mechanical properties and performance characteristics have been achieved by intelligently tuning the printing parameters. Additionally, the apparatus will serve as an invaluable validation suite for high-fidelity models that attempt to simulate the process.ACKNOWLEDGEMENTSThe authors would like to thank The ExOne Company and MicroFab, Inc. for their help and in-kind support during this work. The authors would also like to thank Advanced Powders & Coatings, Inc, (A GE Company) for donating powder samples used in early stages of the project.REFERENCES[1]E. Sachs, M. Cima, P. Williams, D. Brancazio and J. Cornie, "Three Dimensional Printing: Rapid Tooling and Prototypes Directly from a CAD Model," Journal of Engineering for Industry, vol. 114, pp.

Zafar, S. Nadimi and M. Ghadiri,[7]"Jamming during particle spreading in additive manufacturing," Powder Technology, vol. 338, pp. 253-262, 2018.145[9] W. Zhang, A. Mehta, P. Desai and C. F. Higgs III, "Machine learning enabled powder spreading process map for metal additive manufacturing (AM)," in Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium, Austin, TX, 2017.[10] N. B. Crane, J. Wilkes, E. Sachs and S. M. Allen, "Improving accuracy of powdersintering-based SFF processes by metal deposition from nanoparticle dispersion," in Proceedings of the 15th Annual International Solid Freeform Fabrication Symposium, Austin. TX, 2005.[11] Y. Bai and C. B. Williams, "Binder jetting additive manufacturing with a particle-freemetal ink as a binder precursor," Materials & Design, vol. 147, pp. 146-156, 2018.[12] L. O. Grant, M. B. Alameen, R. Carazzone, C. F. Higgs III and Z. C.

147, pp. 146-156, 2018.[12] L. O. Grant, M. B. Alameen, R. Carazzone, C. F. Higgs III and Z. C. Cordero,"Mitigating distortion during sintering of binder jet printed ceramics," in Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium, Austin, TX, 2018.[13] T. Fan, "Droplet-powder impact interaction in three dimensional printing," Ph.D.[14]dissertation, Dept. of Mech. Eng., MIT, Cambridge, MA, 1996. J. Bredt, "Binder stability and powder-binder interaction in three-dimensional printing," Ph.D. dissertation, Dept. of Mech. Eng., MIT, Cambridge, MA, 1995.[15] E. Sachs, M. Cima, J. Cornie, D. Brancazio, J. Bredt, A. Curodeau, T. Fan, S. Khanuja, A. Lauder, J. Lee and S. Michaels, "Three-Dimensional Printing: The Physics and Implications of Additive Manufacturing," CIRP Annals - Manufacturing Technology, vol. 42, no. 1, pp. 257-260, 1993.[16] P. Baker, "Three dimensional printing with fine metal powders," Ph.D. dissertation, Dept.of Mech.

PERIMENTAL INVESTIGATION OF LASER METAL DEPOSITION OF FUNCTONALLY GRADED COPPER AND STEELVinay Kadekar, Sashikanth Prakash and Frank Liou University Of Missouri – Rolla Rolla, Mo 65401AbstractLaser metal deposition is an emerging technology for producing fully dense metallic parts. This process shows a promising future for the deposition of functionally graded steel - copper alloys. Good thermal conductivity of copper and a high wear resistance of steel can be achieved in dies and cores. However, to accomplish this, there are many issues to be resolved, such as the formation of an undesirable phase, solidification cracking, porosity at the interface and difference in thermal coefficient of expansion between steel and copper.

The influences of process variables, such as laser power, laser scan speed, composition, powder flow rate, on the success of the process, should be studied.IntroductionH-13 tool steel is a widely used material for the manufacturing of dies due to its high wear resistance, but the low thermal conductivity limits the melt cooling rate and subsequent mold cycle time. Properties such as high thermal and electrical conductivity of copper make it a suitable material for the die casting industry. Copper is currently being considered for the deposition on dies made out of steel to enhance the thermal conductivity [1]. One of the processes to achieve this would be laser welding of copper onto steel. However, copper has a low absorptivity of the laser light and hence a high reflectivity. These properties of copper makes it difficult for laser welding as most of the laser light is reflected and a small portion of the energy is absorbed.

These properties of copper makes it difficult for laser welding as most of the laser light is reflected and a small portion of the energy is absorbed. Also the difference in thermal coefficient between copper and steel creates a large thermal stress at the clad interface resulting in a failure.Laser Metal Deposition (LMD) is a layered deposition process where the metal powder is focused on a substrate and a laser beam is used to melt the powder and deposit it. The advantage of this process is that, complex geometries can be constructed with high degrees of accuracy to achieve near net shape with a solid model of the part. Figure 1 shows a Laser Aided Material Processing system ( LAMP) at the University Of Missouri- Rolla which consists of a 2.5 KW Nd: YAG Rofin Sinar laser (at TEM00) integrated with a 5-Axis FADAL CNC. The coaxial dual powder feeder system feeds the powder through four tubes which are connected to the nozzle.

The coaxial dual powder feeder system feeds the powder through four tubes which are connected to the nozzle. The advantage of using a dual powder feeding system is that it is capable of delivering two types of powder for functionally graded material parts. The laser deposition system has been integrated with a 5- Axis CNC machine to ensure a high finish quality, as the excess material can be198Figure 1. Laser Aided Material Processing System (LAMP).machined. One more advantage of using the 5- Axis CNC machine is the capability to produce overhang parts. This avoids the use of support material thereby minimizing the build time and the time required for post processing.There are a large range of layer thicknesses as well as deposition rates that can be achieved using laser deposition. The rate of deposition can be increased by increasing the laser power, powder flow rate and by decreasing the traverse speed.

The rate of deposition can be increased by increasing the laser power, powder flow rate and by decreasing the traverse speed. However, the requirement of a good part quality puts a limit on optimal deposition speeds. Both the layer thickness and the volume deposition rates are affected predominately by the specific energy and powder mass flow rate. Considering the factors, it has been reported that there is a positive linear relationship between the layer thickness and adjusted specific energy for each powder mass flow rate [2].Design of ExperimentsExperimental design is a critically important tool in the engineering world for improving the performance of a manufacturing process. The main advantage of this design is that an optimized result can be achieved in the least possible time. Hence, this tool can be implemented in designing a set of experiments for the deposition of copper on H13. The steps that should be used when designing such an experiment are given below.1.

Hence, this tool can be implemented in designing a set of experiments for the deposition of copper on H13. The steps that should be used when designing such an experiment are given below.1. Identifying the potential design factors that the experimenter may wish to vary during the experiment. A total of 7 factors with 2 levels were chosen in this process as shown in Table 1.2. In selecting the response variable, the experimenter should be certain that this variable really provides useful information about the process. The response variables considered under study are porosity, cracks, micro hardness and thermal conductivity.3. The number of experiments to be conducted depends on the number of factors and their levels. A total of 7 factors with 2 levels each would give rise to 128 experiments ( 27 experiments). Hence a 27-3 fractional factorial design can be used199to reduce the number of experiments to 16 runs.

Hence a 27-3 fractional factorial design can be used199to reduce the number of experiments to 16 runs. Table 2 shows the experimental layout to run the 16 experiments with complete randomization.4. Statistical methods should be used to analyze the data so that results and conclusions are objective rather than judgmental in nature.5. Confirmation experiments should be done to validate the accuracy of the experiment.Factors (unit)Levels1. Laser Power ( W) 2. Feed rate ( ipm) 3. Overlap ( %) 4. Outer gas ( psi) 5. Inner gas (psi) 6. Powder flowrate ( g/m) 7. Standoff distance ( in)1 500 20 25 8 3 3 0.3+1 800 40 35 12 5 6 0.4Table 1. Control Factor and their Levels.RunOrder Laser power Feed rate Overlap Outer gas Inner gas Powder flowrate Standoff distance1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 Table 2.

Control Factor and their Levels.RunOrder Laser power Feed rate Overlap Outer gas Inner gas Powder flowrate Standoff distance1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 Table 2. Completely randomized experimental layout.1 1 -1 -1 1 1 -1 -1 1 -1 1 -1 1 1 -1 11 -1 1 1 -1 -1 -1 -1 1 1 -1 -1 1 1 -1 11 -1 -1 1 1 1 -1 -1 1 1 -1 1 1 -1 1 -11 1 -1 -1 -1 1 1 -1 -1 1 -1 1 1 1 -1 -11 1 1 -1 -1 -1 -1 -1 1 -1 1 1 1 -1 1 -11 1 1 -1 1 -1 1 -1 -1 1 -1 -1 1 -1 1 1DiscussionThe basic chemical composition of H13 is shown in Table 3. H13 is a chromium- molybdenum hot working steel, which can be used in either hot or cold working conditions. It has a high resistance to thermal fatigue, high toughness and low thermal conductivity when used conventionally. On the other hand copper has almost 13 times higher thermal conductivity than H13. This high thermal conductivity of copper is potentially significant in increasing the heat conduction in H13 dies.

On the other hand copper has almost 13 times higher thermal conductivity than H13. This high thermal conductivity of copper is potentially significant in increasing the heat conduction in H13 dies. However, copper is typically considered as a contamination is steel. Figure 2 shows an iron copper system with a high solidification temperature range coupled with low solid solubility which2001 -1 -1 1 -1 1 1 -1 -1 -1 1 -1 1 -1 1 1Element Composition (wt %)C 0.4Cr 5.2V 1.0Mo 1.5Si 1.0Mn 0.4Ni 0.3Table 3. Composition of H13 Tool SteelFigure 2. Cu – Fe phase diagram.results in solidification cracking. This resultant cracking is a strong function of copper concentration in the deposit. Cracking is seen extensively when the percentage of copper is between 10-50 wt%, while deposits with copper concentrations below 5 wt% and above 55 wt% show a relatively low cracking [3,4].

Cracking is seen extensively when the percentage of copper is between 10-50 wt%, while deposits with copper concentrations below 5 wt% and above 55 wt% show a relatively low cracking [3,4]. At the intermediate copper concentration, say up to 55 wt%, copper wets in between the dendrites which thereby decrease the ability to withstand solidification stresses, resulting in the formation of cracks [5]. One way of reducing the cracking is to introduce an intermediate Nickel layer between the copper and steel deposit. Nickel is easily soluble in copper and steel, thereby making it a good choice for intermediate layer. The other important factor to be examined is the porosity. This may result due to the copper powder’s dendrite morphology. Further research has to be conducted to closely examine the cracks and the porosity.

This may result due to the copper powder’s dendrite morphology. Further research has to be conducted to closely examine the cracks and the porosity. This can be done by conducting iterative experiments to hone in on the best possible level values of the control parameters that are least sensitive to the variations and are more repeatable.201Fe balReferences1. F.F Noecker II and J.N. DuPont, “ Functionally Graded Copper-Steel Using LENS Process”, Rapid Prototyping of Materials, Symposium Proceedings, Columbus, OH, United States, Oct. 7-10, 2002 (2002), 139-1472. Mazumder J., Choi J., Nagarathnam K., Koch J., and Hetzner D., “The Direct Metal Deposition of H13 Tool Steel for 3-D Components”, JOM, 49(5), 1997, pp. 55-60.3. Vainerman, A.E., Osetnik, A.A. “The Formation of Cracks During the Deposition of Copper Alloys to Steel”, Automatic Welding, Vol 21, June, 1968, pp 22-25.4. Dixon, B.

perimental Investigation of Mechanical Properties of 3D-Printing Built Composite MaterialYi-Tang Kao, Trace Dressen, Dong Sung (Danny) Kim, Sima Ahmadizadyekta, Bruce L. Tai *Department of Mechanical Engineering, Texas A&M University, College Station, 77843Keywords: 3D printing, additive manufacturing, composite materialsAbstractThis paper studies the mechanical behaviors of a new composite material manufactured by 3D printing and polymer impregnation techniques. This composite uses 3D-printed plaster with an open-cellular structure as a frame to encapsulate the silicone resin (PDMS) to form a solid body. Because of the vastly different characteristics of the materials that make it up, the composite could have a wide variety of mechanical behaviors. In this study, design of experiment was performed with four-point bending tests using different composition ratios and sizes of open cells to determine the mechanical properties of the composite.

In this study, design of experiment was performed with four-point bending tests using different composition ratios and sizes of open cells to determine the mechanical properties of the composite. These properties include maximum flexural stress (σmax), flexural secant modulus of elasticity (Ef), and toughness indices (I5 and I20). The experimental results show that both Ef and σmax are proportional to the plaster content and the unit cell size, while I20 had an opposite trend. The Ef ranged from 20 to 280 MPa, and σmax ranged from 0.3 to 1.2 MPa for a 25%-75% plaster content and 3.25-6.5 mm cell size. Statistical analysis further confirmed the differences between these cases.

Statistical analysis further confirmed the differences between these cases. This paper has demonstrated the capability of this composite to exert different mechanical properties for functional applications.IntroductionAdditive manufacturing (sometimes referred to as “3D printing”) is an emerging technology used to build three-dimensional structures based on layer-by-layer deposition [1]. The main advantages of this technology include the ability to create almost any complex and light structure [2], minimize the material used, and decrease the cost for manufacturing small amounts of parts. The revolution of 3D-printing technology is undeniable, and it is evolving toward producing functional materials for practical use, such as composites. Recently, some 3D printers have been built with the ability to print multiple materials at the same time.

Recently, some 3D printers have been built with the ability to print multiple materials at the same time. For example, a commercially available printer, ProJet 5500X (3D Systems, Rock Hill, SC), prints rigid plastic and rubber-like materials with different material transparency and colors. The Objet Connex500 (Stratasys, Edina, MN) can print photopolymer, digital ABS, and rubber-like materials with varying material transparency and colors. Thus far, the amount of the available printing materials is still limited, and there is no significant difference in the properties of the materials used to create composites.Other research on additively fabricating composites also continues. Laminated object manufacturing builds laminated fiber composites such as ceramic matrix composites (SiC/SiC) and polymer matrix composites (glass/epoxy) [3,4].

Laminated object manufacturing builds laminated fiber composites such as ceramic matrix composites (SiC/SiC) and polymer matrix composites (glass/epoxy) [3,4]. A selective laser sintering/melting technique has also been used for manufacturing various composite materials, such as metal-metal, polymer-metal, polymer-ceramic, and metal-ceramic composites [5]. H. Kyogoku et al. reported the applicability of laser melting for manufacturing Ti-Ni shape memory alloys [6]. D.D. Gu et al. studied mechanisms in laser melting for a Cu–CuSn–CuP mixed powder [7]. K.K.B. Hon and T.J. Gill presented an experimental study in selective laser sintering for Silicon Carbide/Polyamide matrix composites [8]. H.S. Chung and Suman Das investigated selective laser sintering for functionally graded materials, which are composites of Nylon-11 and different volumes of glass beads [9]. These studies provide a variety of options in material preparation to strengthen 3D- printed parts.

These studies provide a variety of options in material preparation to strengthen 3D- printed parts. However, such composite printing focuses on the material itself and does not have the flexibility to create a part with structurally anisotropic properties for directional or localized strengthening. This paper, therefore, presents a new concept for the future application of 3D printing technology that can construct a composite of not only multiple materials but also with selected structural strength. The proposed composite consists of two or more vastly different materials and thus possesses a broader range of mechanical properties (from brittle to ductile) that could be tailored by manufacturers. Furthermore, different geometrical arrangements of these two phases can create different directional strengths. This concept is similar to the reinforced concrete used in construction with concrete as a rigid base and rebar as a tough addition.

This concept is similar to the reinforced concrete used in construction with concrete as a rigid base and rebar as a tough addition. Owing to the flexibility of 3D printing, the desired structure design can be easily fabricated. In this study, a conceptual prototype was made with a powder-bed printer for the brittle phase and then combined with a silicone material as the ductile phase. The objective is to experimentally measure the changes in mechanical properties of the built composite and determine the effects of composition and structure.In this paper, Section 2 will detail the material and methods used in the experimentation, followed by the results and discussion in Section 3. Section 4 presents the major conclusions and future works of this topic.Materials and MethodsSample preparationA four-point bending test was utilized to examine the material properties of the composites.

The test sample was designed to be 128 mm × 13 mm × 6.5 mm based on an open-cellular unit cubic structure to accommodate the silicone material, as shown in Fig. 1. The porosity of the part is determined by a unit cubic cell with a specific void-to-body ratio as shown in Fig.1 (a). The unit cells are assembled to form the part as shown in Fig. 1 (b) and (c). The complementary percentage of this ratio is defined as the composition ratio. A different structure could be made with an identical composition ratio. For example, the sample in Fig. 1 (b) is made of unit cells 6.5 mm in outer length with a composition ratio of 50%, compared to the sample in Fig. 1 (c) with 3.25 mm long cells and the same composition ratio.(b)(a)(c)Figure 1 Sample design: (a) a unit cubic cell, (b) the frames based on the unit cubic cells of 6.5 mm and (c) 3.25 mm.ProJet 160 (3D Systems, Rock Hill, SC), a commercial powder-bed 3D printer, was selected in this study for creating the frames.

The powder material for this printer was VisiJet PXL, which contained 80 to 90% calcium sulfate hemihydrate (also known as plaster). A low elastic modulus silicone resin (PDMS), Sylgard 184 (Dow Corning, Midland, MI), was used along with the plaster frame to build the composite material. A mold, as shown in Fig. 2, was built using three separate aluminum plates to hold and cure the composite.Figure 2 The mold for fabricating the composite samplesThe steps for fabricating composite samples are summarized as follows: First, the silicone resin was fully mixed with hardener (Sylgard 184 curing agent, Dow Corning) in a disposable plastic cup with a 10:1 ratio according to the material datasheet. Then, the mixed silicone resin was completely degassed and poured into the mold. The pure plaster frames were baked for 10 minutes at 80˚C in an air oven and then placed into the slot of the mold filled with the silicone resin.

The pure plaster frames were baked for 10 minutes at 80˚C in an air oven and then placed into the slot of the mold filled with the silicone resin. The uncured composites were degassed again to remove any remaining air bubbles. Finally, the whole mold was placed into the air oven to bake at 80˚C for 2.5 hours to fully cure the composite. Samples were then removed from the mold for testing.Experimental setup and designThe experimental setup for the four-point bending test is shown in Fig. 3, consisting of a linear actuator and a force transducer. The linear actuator (L70, Moog Animatics, Milpitas, CA) was driven by a servo-motor for a precise position and accurate feed rate control. Due to an anticipated small force, a high-sensitivity and high-response frequency force dynamometer (Model 9272, Kistler, Winterthur, Switzerland) was used to capture the force data.

Due to an anticipated small force, a high-sensitivity and high-response frequency force dynamometer (Model 9272, Kistler, Winterthur, Switzerland) was used to capture the force data. An amplifier, a shielded connector block, and a data acquisition device (PCle-6321, National Instruments, Austin, TX) were used for data collection, along with LabVIEW as a data recorder.As shown in Fig. 3, the loading nose and the loading support both had a cylindrical contact surface with a 5 mm radius. The support span was 128 mm and the loading span was 64 mm, in accordance with the ASTM-D7264 standard [10].

The support span was 128 mm and the loading span was 64 mm, in accordance with the ASTM-D7264 standard [10]. The feed rate of the loading support was 1.27 mm/min (0.05 inches/min); the maximum displacement of the loading noses was 50.8 mm (2 inches), which was the hardware limit with this setup.Linear actuatorLoading nose64 mmSampl e128 mmLoading supportAmplifierDynamometerFigure 3 Experimental setup configured for the four-point bending testIn design of experiment, four different composition ratios were made for testing, including 25%, 50%, 75%, and 100%. The results with these different composition ratios were compared to the results of pure silicone and pure plaster samples to quantify the strengthening effects of the composites. In addition, for every composition ratio of a composite, there were two different structures: 6.5 mm and 3.25 mm long unit cells.

In addition, for every composition ratio of a composite, there were two different structures: 6.5 mm and 3.25 mm long unit cells. These different structures were used to determine how the structures affected the material properties beyond the composition ratio. In total, there were six cases and one pure plaster case. Samples are shown in Fig. 4, where large and small unit cells are denoted by L and S, respectively. Each sample had five replicas for the bending test.Figure 4 Completed composite samples with four composition ratios (25%, 50%, 75% and 100%) and two types of unit cells (L and S)Three mechanical properties were obtained from the four-point bending test, including maximum flexural stress (σmax), flexural secant modulus of elasticity (Ef), and toughness indices (I5 and I20). Following the ASTM standard, σmax is calculated by Eq.

Following the ASTM standard, σmax is calculated by Eq. (1)2max43bhPLwhere P is the maximum force applied, b is the width of sample (13mm), h is the height of sample (6.5 mm), and L is the support span (128 mm) of the four point bending setup. The Ef is calculated by Eq. (2)3317.0bhmLEfwhere L, b, and h are the same as those in Eq. (1), and m is the slope of the force-deflection curve within the elastic (linear) region. The toughness is represented by toughness indices (I5 and I20) described in ASTM 1018 [11]. Since the four-point bending samples do not have a notch at the mid-span, the flexural toughness could not be obtained. Instead, the toughness indices, based on the force-deflection curve, are described by11201155.10 ,)()3(AAIAAIwhere A is the area under the curve at a certain deflection and δ1 is the first cracking point.

The first cracking point is defined by the transition from a linear region into a non-linear region, equivalent to the yielding point in a tensile test.Experimental ResultsBoth qualitative and quantitative comparisons are conducted in this section. Fig. 5 shows the average force-deflection curves of the composite materials to illustrate the differences in their mechanical behaviors. Each curve representing a composite sample is calculated based on the test results of five replicas. The pure plaster sample is not shown in the figure as it has a maximum average force of 36.5 N, with a rupture deflection of 2.9 mm, both of which are vastly different than the composite samples. This indicates that the pure plaster is relatively brittle. As shown in Fig. 5, 75% plaster samples (both L and S) break before the maximum deflection of 50.8 mm since the plaster structure dominates the mechanical behavior.

As shown in Fig. 5, 75% plaster samples (both L and S) break before the maximum deflection of 50.8 mm since the plaster structure dominates the mechanical behavior. However, in the cases of the 50% and 25% plaster samples, no rupture occurs prior to the maximum deflection.For the elastic performance, all the samples display a nearly elastic behavior when the deflections of the midpoints are lower than 3.5 mm, regardless of the composition ratios and structures. This is because the elastic region is, again, dominated by the plaster structure. Moreover, this is also the reason for crack occurrences at the same strain (midpoint deflection). In general, the slopes of the force-deflection curves significantly decrease once the deflections exceed 3.5 mm. At this point, the plaster frame of a sample cracks and the silicone portion begins to act as rebar to hold the structure together and prevent crack propagation.

At this point, the plaster frame of a sample cracks and the silicone portion begins to act as rebar to hold the structure together and prevent crack propagation. Therefore, 3.5 mm is defined as the first cracking point (δ1) for the following toughness indices calculation. After the δ1, the force either increases slowly or remains at a nearly constant level. This is because of the exceptional ductility of the silicone resin.(1)(2)(3)Figure 5 Averaged force-deflection curves for the composite samples.Based on the averaged data, toughness indices of the composite samples were calculated and are shown in Fig. 6. All the samples have similar I5. This indicates that, immediately after the δ1, the force-deflection curves change in a similar trend for all the samples despite different magnitudes. In contrast, the results of the I20 are quite different for all the samples.

In contrast, the results of the I20 are quite different for all the samples. Samples with a lower amount of the plaster (i.e., lower composition ratio) possess higher I20, meaning the silicone resin contributes to the toughness more at a larger deflection. In addition to the composition ratio, the sample structure (i.e., unit cell size) plays an important role in the toughness index I20. The I20 of large-cell samples with 25% and 50% composition ratios are smaller than those with small cells. However, the 75% samples display an opposite phenomenon.Figure 6 Toughness indices of plaster-silicone composite samples.Fig. 7 shows the statistical results of flexural secant moduli of elasticity (Ef) for all composite samples. The slope was fitted in the first 0.04 mm of the force-deflection curve for each sample to find Ef. Pure plaster and pure silicone cases are not shown in the figure due to their relatively extreme properties.

Pure plaster and pure silicone cases are not shown in the figure due to their relatively extreme properties. The pure plaster solid sample has an Ef up to 1611 MPa, while the Ef of a pure silicone resin is nearly zero because it could not resist any bending moment. Resultsin Fig. 7 suggest an increasing Ef with the composition ratio. The Ef is dominated by the amount of plaster in the structure because it possesses a much higher stiffness than silicone resin. Statistical analysis using a t-test also confirms a significant difference in Ef when the composition ratio increases. On the other hand, for a given composition ratio, samples with a larger unit cell structure tend to have a higher Ef. Statistically, there is a significant difference between small- and large- cell samples, excepting the case of 75% due to the variations.Figure 7 Flexural secant moduli of elasticity (Ef) for plaster-silicone composite samples.

* represents a statistical significance (p < 0.05) between different composition ratios; + means a statistical difference between small and large cell samples.Fig. 8 shows the maximum stress (σmax) for all the samples. The σmax of the pure plaster samples was 6.15 MPa on average, and that of the pure silicone was not available due to a nearly-zero stiffness in the bending test. For the composite materials, σmax increases with the composition ratio because the plaster material is the major source of the strength. Statistical results revealed that the samples with the same unit cell but different composition ratios could be different in σmax, particularly for small-cell samples. If comparing σmax based on the sample structure, the samples with a larger unit cell have higher σmax than those with small unit cells. This phenomenon is similar to the results of Ef in Fig. 7.Figure 8 Maximum stress (σmax) for plaster-silicone composite samples.

This phenomenon is similar to the results of Ef in Fig. 7.Figure 8 Maximum stress (σmax) for plaster-silicone composite samples. * represents a statistical significance (p < 0.05) between different composition ratios; + means a statistical difference between small and large cell samples.DiscussionThe results suggested that both composition ratio and structure can affect the mechanical properties of the plaster-silicone composites. Since the data was obtained via four-point bending tests, the flexural secant moduli of elasticity (Ef), maximum stress (σmax), and toughness index (I20) are largely dependent on the resistance against the moment. That is, a resultant bending stiffness (i.e., the production of elastic modulus and moment of inertia) plays an important role in the composite materials.In regards to Ef (in Fig.

That is, a resultant bending stiffness (i.e., the production of elastic modulus and moment of inertia) plays an important role in the composite materials.In regards to Ef (in Fig. 7), the samples with large and small unit cells are anticipated to have the same elasticity under a uniaxial tension; however, samples with a larger unit cell display higher moduli in a bending case due to a higher moment of inertia. The difference in the 75% case could be due to variations in the samples.Similarly, in regards to σmax (Fig. 8), structures with larger cells also possess higher σmax. It is important to note that this stress is calculated based on a homogenous cross-section. Therefore, this stress does not mean an absolute higher stress on the outer surface of the structure. Instead, it is a result of a higher Ef under the same strain (deflection). Another possible reason is that a larger unit cell has a better structural integrity than a smaller unit cell.

Instead, it is a result of a higher Ef under the same strain (deflection). Another possible reason is that a larger unit cell has a better structural integrity than a smaller unit cell. The powder printing material has a size of 10 to 100 μm in diameter. Thus, with a feature size in mm or sub-mm scales, the number of particles within each structural feature could be critical to its strength.Having seen the differences in the mechanical properties of these samples, it would be of interest to see if there exists a model to predict these properties for design purposes. A typical model for composite materials is known as the rule of mixtures. This model gives the upper and lower bounds of a composite material with the reinforced fibers parallel or perpendicular to the stress flow, known as the iso-strain and iso-stress conditions. The equations are given by Eq. (4) and Eq.

The equations are given by Eq. (4) and Eq. (5), respectively [12].ppsscEVEVEpsspspcVEVEEEEwhere, in our case, Ep and Es are the elastic moduli of plaster and silicone resin, and Vp and Vs are their volume fractions, respectively. The Es is set as 1.78 MPa, adopted from the literature [12], since it is too low to be measured in the four-point bending test. The EP in this study is 1.61 GPa. Fig. 9 shows the upper and lower bounds of this type of composite material. As expected, both large and small cell samples fall within the boundaries. An empirical rule of mixture can be established when a sufficient amount of data points are added to the plot. In addition, it can also be found that the small-cell samples are closer to the average of the upper and lower bounds, which is an isotropic mixing condition.

This fact verifies that small-cell samples act more like a homogeneously mixed material due to the fine grid inside the material.(4)(5)Figure 8 The upper and lower bounds of the composite samples based on the rule of mixtures.Although the current data and statistical analyses have suggested the basic trends of properties change in accordance to the composition ratio and structure, there were still limitations in this study: first, the variations were large, which resulted in an insufficient statistical power to distinguish samples; second, more levels in both structure and composition ratio in the design of experiments are needed to lead to more solid conclusions; third, only one type of silicon resin was used, so it is unknown if the results are applicable to general cases.ConclusionsThis paper presented a new concept of composite material built using 3D printed part as the brittle material and durable silicone material as the ductile material, aiming to create a reinforced structure for functional applications.

Prototype samples made of 3D printed plaster and silicone resin were fabricated with various composition ratios and structures for the four-point bending test in order to obtain the mechanical properties. Based on the data, the major findings in this paper are:Given the extreme properties of brittle and ductile materials, a composite with a wide variety of mechanical properties can be created.The brittle material dominates the strength and elastic modulus of the material, while the ductile material controls the toughness.

A balance between these two materials could maximize the material’s functionality.The mechanical properties of the composites are determined by both the composition ratios of these two materials and the structural configuration (i.e., unit cell size).The composites composed of small unit cells tend to behave as a homogenous mixture of two materials, but a better strength and stiffness are often provided by larger unit cells, particularly in bending.The rule of mixtures could be applicable to this type of composite.The future works of this study include testing different combinations of the materials to generalize the rule of mixtures model, investigating the failure mechanism of the composites, and exploring a hybrid manufacturing process to build the composites at a time without manual operations in molding.AcknowledgmentsThis research is supported by the National Science Foundation grant #1522877.

Theauthors would also like to acknowledge the material supports from Dow Corning.References1.2.3.4.5.6.7.8.9.10.11.Hung, K.C., Tseng, C.S., and Hsu, S.H., Synthesis and 3D printing of biodegradable polyurethane elastomer by a water-based process for cartilage tissue engineering applications. Adv Healthc Mater, 2014. 3(10): p. 1578-87. Wong, K.V. and Hernandez, A., A Review of Additive Manufacturing. ISRN Mechanical Engineering, 2012. 2012: p. 1-10. Kruth, J.P., Leu, M.C., and Nakagawa, T., Progress in Additive Manufacturing and Rapid Prototyping. CIRP Annals - Manufacturing Technology, 1998. 47(2): p. 525-540. Donald A. Klosterman, Richard P. Chartoff, Brian Priore, Nora Osborne, George Graves, Allan Lightman, Sung S. Pak, and Weaver, J., Structural Composites via Laminated Object Manufacturing (LOM), in Solid Freeform Fabrication Symposium. 1996. p. 105- 116.

PERIMENTAL INVESTIGATION OF PROCESS PARAMETERS ON 64 SLICE SPIRAL CT SCANNER OF MEDICAL MODELSY. Ravi Kumar, A. Manmadhachary and L. KrishnanandDepartment of Mechanical Engineering, National Institute of Technology Warangal, Telangana, 506004, India.AbstractRapid Prototyping (RP) is one of the advanced manufacturing methods to develop medical models. These models are generated by 3-Dimensional (3D) Computer Aided Design (CAD) model using Computed Tomography (CT) images. One of the advanced CT scanners to capture the large volume of tissues in shorter scan time is 64 slice spiral CT scanner. While developing these medical models, dimensional and volumetric errors occur due to Beam Hardening (BH) effect. This work has led to explore the influence of various CT Image acquisition parameters on the dimensional and volumetric errors, which are evaluated experimentally.

This work has led to explore the influence of various CT Image acquisition parameters on the dimensional and volumetric errors, which are evaluated experimentally. A L9 orthogonal array and signal to noise ratio are applied to study performance characteristics of CT image acquisition parameters like tube voltage, tube current and pitch. The experimental results are analyzed by using the analysis of variance (ANOVA) method and significant factors are identified. In this work, it has been concluded that there is a reduction of dimensional error from 1.43 mm to 0.52 mm and volumetric error from 6793 mm3 to 3892 mm3.Key words: CT, Dimensional error, Volumetric error, Rapid prototyping, CT image acquisition, ANOVA.IntroductionRapid Prototyping (RP) is the automatic construction of physical objects using solid freeform fabrication.

In other words, RP is an additive fabrication process refers to a group of technologies used for building physical models, prototypes, tooling components, finished series production parts all from 3-Dimensional (3D) Computer Aided Design (CAD) data, Computed Tomography (CT) or Magnetic Resonance Imaging (MRI) scans, or data from 3D scanning systems [1]. Unlike machining processes, which are subtractive in nature, RP’s additive systems join together liquid, powder, or sheet materials to form parts [2]. Parts that may be difficult or even impossible to manufacture by any other method can be produced by additive systems [3]. Based on thin, horizontal cross sections taken from a 3D computer model, they produce plastic, metal, ceramic, or composite parts, layer upon layer.Most of these RP systems generally adopt the standard processing steps to produce a part.

They are created from 3D CAD modelling of the part to be produced, convert the CAD files into a RP industrial standard data file format, import the data file into the RP program of that system, check for errors and missing data in the data file, perform corrections where necessary, digitally slice the 3D CAD solid model into horizontal layers, send the sliced data to the RP machine for production and post-process the prototype such as removing redundant materials [2].RP is one of the advanced manufacturing techniques to produce 3D physical medical models. These medical models are widely used for several applications like visualization,diagnosis, surgery planning [4], design of implants, external prosthesis, surgical templates, production of artificial organs, communication between the medical team and/or medical doctors and patients, and teaching aids [5]. In the present work, CT images are considered as input data to produce 3D physical medical models.

In the present work, CT images are considered as input data to produce 3D physical medical models. During acquisition of the CT images from the human anatomy, usually various dimensional and volumetric errors occur due to Beam Hardening (BH) effect [6].The presence of dense material in the middle portion of a uniform cylindrical phantom influences the X-rays in becoming hard as compared to the rays passing through the edges, which has shown in Fig.1. The hardened beam has less attenuation and intense as it reaches the detector; therefore, the difference in the attenuation profiles of this from that of ideal with beam hardening artifact. This artifact is also called “cupping” artifact because the hardening is more in the center and less on the periphery, it resembles a cup, which has shown in Fig. 1.Fig. 1: BH effect on cylindrical phantomDual energy CT systems scan at two energy levels, which enable BH correction [7].

1.Fig. 1: BH effect on cylindrical phantomDual energy CT systems scan at two energy levels, which enable BH correction [7]. In dual energy CT systems, the X-ray beam energy is half the single X-ray CT scan systems so as to reduce the radiation dosage to the patient [8]. But this lower energy X-ray beam gives raise to image noise (photon starvation). So instead of dual energy CT scanner, single X-ray beam CT scanners are being used by manipulating tube voltage, tube current and pitch. To overcome BH effect, increasing in tube voltage [9] and tube current [10] ultimately increases the X-ray beam energy. Selections of optimum scanner parameters like tube voltage, tube current and pitch are the very important factors in avoiding dimensional and volumetric errors.

Selections of optimum scanner parameters like tube voltage, tube current and pitch are the very important factors in avoiding dimensional and volumetric errors. Manufacturers and many authors are trying to eliminate the dimensional and volumetric errors using different cylindrical phantoms in a range of different sizes [11], but patient anatomy never exactly matches these cylindrical phantoms. X-ray beam photon attenuation is proportional to the density of the phantom [12]. In clinical practice generally bone is a phantom, but acrylic type of materials is used in experimental analysis. The acrylic materials and bone having different densities and Hounsfield Unit (HU) [11] values, due to this reason the variation will be observed in practice. In this study, a human dry mandible has been considered as a phantom for experimentation.Experimental SetupAn adult dry mandible (Fig. 2) as a phantom for CT image acquisitions was used in this study to find dimensional and volumetric errors.

2) as a phantom for CT image acquisitions was used in this study to find dimensional and volumetric errors. This dry mandible has 110 mm length (X-axis), 80 mm width (Y-axis) and 65 mm height (Z-axis).Fig. 2: Dry mandible phantomThe phantom was scanned with a 64 slice spiral CT scanner (Light Speed VCT, GE Medical Systems). The primary scan was done with default setting parameters of tube voltage 120 kV, tube current 300 mA and pitch 0.516 [13]. However, the other scans were done with different tube voltages, tube currents and pitches (shown in the Table 1) using L9 orthogonal array.

However, the other scans were done with different tube voltages, tube currents and pitches (shown in the Table 1) using L9 orthogonal array. Further, optimized parameters produced were used for the last scan is a conformation test.Table 1 CT Scanner acquisition parametersParameter Tube voltage (kV) Tube current (mA) Pitch (table movement/slice thickness)Level of Parameter 2 100 400 0.9841 120 300 0.5163 80 500 1.375Scanning of the phantom with three different parameters (tube voltage, tube current and pitch) having three levels requires 33 (27) experiments. Taguchi method [15] uses a special design of orthogonal array to conduct the experiments and gives the optimize parameters. The degrees of freedom for three parameters in each of three levels were calculated by following equation.Degree of Freedom (DOF) = number of levels -1--------------------------------------- (1)In this study three parameters and three levels were considered.

So here L9 standard orthogonal array is selected. This L9 orthogonal array has eight DOF, in which 6 were assigned to three factors (each one 2 DOF) and 2 DOF was assigned to the error. Based on this L9 orthogonal array (shown in the Table 2) nine scans were done with different tube voltage, tube current and pitch. The images in Digital Imaging and Communication in Medicine (DICOM) file format is used for developing 3D CAD model by utilizing Materialize Interactive Medical Image Control System (MIMICS) software (version 14.12, Materialise NV, Leuven, Belgium) [14].Table 2 L9 orthogonal array for experimentation Pitch S.No ExperimentTube Voltage (kV) 120 120 120 100 100 100 80 80 80Tube Current (mA) 300 400 500 300 400 500 300 400 500Number1 2 3 4 5 6 7 8 9Exp. No. 1 Exp. No. 2 Exp. No. 3 Exp. No. 4 Exp. No. 5 Exp. No. 6 Exp. No. 7 Exp. No. 8 Exp. No. 90.516 0.984 1.375 0.984 1.375 0.516 1.375 0.516 0.984To find the dimensional error, two X-axis (Fig.

No. 1 Exp. No. 2 Exp. No. 3 Exp. No. 4 Exp. No. 5 Exp. No. 6 Exp. No. 7 Exp. No. 8 Exp. No. 90.516 0.984 1.375 0.984 1.375 0.516 1.375 0.516 0.984To find the dimensional error, two X-axis (Fig. 3a), two Y-axis (Fig. 3b) and two Z-axis (Fig. 3c) dimensions were selected for measurement of the linear dimensions. These axes were similar to the CT scan machine axes (X, Y, and Z). The details of measurements of linear dimensions are shown in Appendix. In this paper the difference between dry mandible linear dimensions to the 3D CAD mandible linear dimensions values are considered as a dimensional error. This can be written as the following mathematical relations:Dimensional error = Dry mandible dimension - 3D CAD mandible dimension------------- (2)(a)(b)(c)Fig. 3: Anatomic landmarks for measurements in (a) X-axis,(b) Y-axis and (c) Z-axisThe 3D CAD mandible linear dimensions (Fig.

3: Anatomic landmarks for measurements in (a) X-axis,(b) Y-axis and (c) Z-axisThe 3D CAD mandible linear dimensions (Fig. 4a) are measured using MIMICS software whereas a digital electronic caliper with least count of 0.01mm is used to measure the dry mandible linear dimensions (Fig. 4b). The linear dimensions were measured three times, for consistency and closest repeated value was chosen as accuracy criteria. Dimensional errors are calculated by Equation 2. These 3D CAD model dimensions and dimensional errors are shown in Table 3.(a)(b)Fig. 4: Dimensions are measured by (a) MIMICSsoftware and (b) digital electronic caliperTable 3 Dimensional measurements of dry mandible and CAD modelsDimensionsX-axis DimensionsY-axis DimensionsZ-axis DimensionsDry mandibleExp. No. 1Exp. No. 2Exp. No. 3Exp. No. 4Exp. No. 5Exp. No. 6Exp. No. 7Exp. No. 8Exp. No.

Dry mandible volume was measured by using water displacement method. In this method, the dry mandible volume is measured by calculating how much water it displaces, or pushes aside when it's placedinto distilled water. Dry mandible volume is calculated by subtracting the volume of the water without the dry mandible from the new measurement with the dry mandible. The measurements were measured in cubic millimeters (mm3). These volumes are shown in Table 4. In this paper the difference between dry mandible volumes to the 3D CAD mandible volume values are considered as a volumetric error. This can be written as the following mathematical relations:Volumetric error = Dry mandible volume - 3D CAD mandible volume --------------------- (3)As per Equation 3, the volumetric errors were calculated and these values are shown in Table 4.Table 4 Volume measurement of dry mandible and CAD modelVolume taken from Dry mandibleExp. No. 1Exp. No. 2Exp. No. 3Exp. No. 4Exp. No. 5Exp.

No. 1Exp. No. 2Exp. No. 3Exp. No. 4Exp. No. 5Exp. No. 6Exp. No. 7Exp. No. 8Exp. No. 9CAD Model Volumetric Error CAD Model Volumetric Error CAD Model Volumetric Error CAD Model Volumetric Error CAD Model Volumetric Error CAD Model Volumetric Error CAD Model Volumetric Error CAD Model Volumetric Error CAD Model Volumetric ErrorVolume (mm3) 32340 25547 6793 25779 6561 25969 6371 26350 5990 26592 5748 26876 5464 27460 4880 27920 4420 28448 3892Taguchi methods [15] have been widely utilized in engineering analysis and consist of a plan of experiments with the objective of experimentation data in a controlled way. In this work Taguchi technique was used to find the optimized parameters of tube voltage, tube current and pitch, with multiple performance characteristics of dimensional and volumetric errors. The experimental results were analyzed by analysis of variance (ANOVA) and significant main factors are identified [16].

The experimental results were analyzed by analysis of variance (ANOVA) and significant main factors are identified [16]. Multiple regression equations are formulated for estimating predicted values of the dimensional and volumetric errors.ResultsIn this method, the average value of the two dimensional error values was considered as an each individual axis of the dimensional error. The average value of X, Y and Z axes were considered as the overall dimensional error for each set of experimentation. The experiment was done with the default setting parameters of tube voltage 120 kV, tube current 300 mA and pitch0.516 and resulted average dimensional and volumetric errors are 1.43 mm and 6793 mm3 respectively, these values are shown in experiment number 1 of Table 5.The dimensional and volumetric error of each set of L9 orthogonal array experimentation values are depicted in Table 5. In this study smaller-the-better quality characteristic is used.

These values are shown in last column of Table 5. Mean of S/N ratio for each level of CT image acquisition parameters were calculated. These are shown in Table 6. In order to analyses the effect of CT image acquisition parameters on the average dimensional and volumetric errors, a main effects plot for S/N ratios of optimized parameters were generated by using Minitab software, this is shown in Fig. 5. From these it was found that the optimal CT image acquisition parameters are tube voltage 80 kV, tube current 500 mA and pitch 0.984.Table 6 Response table for S/N ratios of CT image acquisitionParameter Tube voltage (kV) Tube current (mA) PitchLevels of Parameters 2 1 3 -72.15 -73.34 -71.80 -72.31 -71.55* -71.76Delta -69.82* 3.53 -71.20* 1.11 0.45 -72.00Rank 1 2 3optimized level of parameters(4)70-71-72-73 1.3750.9840.51670-71-72-73 PitchMain Effects Plot for SN ratiosData MeansSignal-to-noise: Smaller is better12010080Fig.

5: Effect of process parameters on dimenstional and volumetric errorsTable 7 shows the results of ANOVA for the dimensional and volumetric errors. This analysis was carried out for a level of significance of 2%, i.e., for 98% level of confidence. The purpose of ANOVA is to investigate, which the CT image acquisition parameter significantly affects the performance characteristics. The factor of tube voltage has 91% of contribution, is the most significant control parameter for CT image acquisition stage.Table 7 ANOVA results of the CT image acquisition parametersParametersTube Voltage Tube Current Pitch Error TotalDegree of freedom 2 2 2 2 8Sum of squares 1810192 156348 12983 9977 1989499Mean square 905096 181.44 15.67 78174 6491 1.30 4988F-ratioPercent contribution 91% 7.8% 0.7% 0.5% 100%A multiple linear regression models are developed in order to predict the values of dimensional and volumetric errors of the medical models.

The developed models are reasonably accurate and can be used for prediction within limits. The regression equations for the dimensional and volumetric errors were generated with the help of Minitab software is as:Dimensional Error = -0.358 + 0.0177 Tube Voltage - 0.00117 Tube Current - 0.0092 Pitch---- (5)Volumetric Error = 1305 + 54.4 Tube Voltage - 3.23 Tube Current + 115 Pitch ----------------- (6)The confirmation experiment is a crucial step and is highly recommended by Taguchi to verify the experimental results. Based on S/N ratio values, optimum process parameters were estimated by using response table and response graph as shown in Fig.5. These optimal CT image acquisition parameters are selected for the confirmation test.

The volumetric error decreased by 2901 mm3. Even after optimization of CT scanner small dimensional and volumetric errors do occur. The elimination of errors occurring during the CT image construction stage is also very essential because it is one of the stages of RP model fabrication.Appendix: Linear Measurements of MandibleBicondylar width (BCoW) - distance between right and left uppermost point of the condyles. Bicoronoid width (BCrW) - distance between right and left uppermost point of the coronoids. Posterior height of right ramus (PHRR) - distance from uppermost point of the head of the condyle to lowermost and posterior point of the mandible angle, right side.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEXPERIMENTAL INVESTIGATIONS OF INHOMOGENEOUS COMPONENT PROPERTIES IN LASER-BASED ADDITIVE MANUFACTURING OF AlSi10MgS. Czink*, V. Schulze* and S. Dietrich*Institute of Applied Materials, Karlsruhe Institute of Technology (KIT), Karlsruhe, GermanyAbstractIn the laser-based additive manufacturing (PBF-LB) process of AlSi10Mg components, the layer-by-layer deposition leads to microscopic and macroscopic thermal effects (shrinkage, residual stresses, overheating) depending on the component geometry. This results in a strong dependence of the microstructure and therefore of the process-induced material properties on the shape of the manufactured component. To analyze this behavior, components with different geometric aspects, such as various construction angles, were built using the PBF-LB process.

To analyze this behavior, components with different geometric aspects, such as various construction angles, were built using the PBF-LB process. In order to perform a spatially-resolved characterization and evaluation of the mechanical behavior of the component, small-scale tensile specimens in the sub-millimeter range were manufactured from representative areas of the components. With the obtained results, design approaches based on local material data can be improved significantly.IntroductionDue to the high degree of individualization and design flexibility, additive manufacturing techniques, such as the laser based powder bed fusion (PBF-LB), have developed from rapid prototyping to established rapid manufacturing processes. In the PBF-LB process, the components are built incrementally by local melting of metal powder with a focused laser beam.

In the PBF-LB process, the components are built incrementally by local melting of metal powder with a focused laser beam. This enables the manufacturing of geometrically complex parts, such as lattice structures or topology optimized components in a single step. One of the most common alloys for the PBF-LB process is AlSi10Mg. Its composition close to the eutectic of the ternary system Al-Si-Mg leads to a good castability in comparison to other aluminum alloys. Due to the combination of good mechanical properties with a low density, the alloy is suitable for applications in lightweight design. Additionally, PBF-LB parts out of AlSi10Mg show similar or even superior mechanical properties compared to traditional casting due to the unique microstructure of the PBF-LB parts. The microstructure is coupled to the high cooling rates in the process of up to 106 K/s [1].

The microstructure is coupled to the high cooling rates in the process of up to 106 K/s [1]. This results in an ultrafine cellular microstructure with primary crystallized Al and a fibrous eutectic Si network at the cell boundaries [2, 3]. However, different processing parameters can influence the cooling conditions and therefore the resulting microstructure [4]. Several studies additionally investigated the influence of the component geometry on the resulting microstructure. For example, Dong et al. studied the change of microstructure for cylindrical AlSi10Mg samples of different sizes [5]. Liu investigated the microstructure and hardness for lattice structures of AlSi10Mg [6]. The study showed that due to the isolating effect of surrounding powder particles at downskin surfaces of overhanging samples the cooling behavior is influenced locally, which leads to a gradient regarding the microstructure. Yue et al.

Yue et al. analyzed this effect with a finite element model and found that the melt pools have a higher size and temperature towards the downskin surface of overhanging structures [7]. In addition to the microstructure, the formation of defects can also be influenced by448the component geometry. Englert et al. showed an increase in keyhole porosity at downskin surfaces at overhanging structures, especially during processing at lower scanning speeds [8]. Within this work, the effect of an inhomogeneous microstructure on the mechanical properties of AlSi10Mg parts were studied. Therefore, representative specimens of different sizes were built with the PBF-LB process under different construction angles. The effects of the microstructure on the local mechanical behavior were evaluated using small-scale tensile bars prepared from the representative specimens.

The effects of the microstructure on the local mechanical behavior were evaluated using small-scale tensile bars prepared from the representative specimens. To analyze the influence of the locally inhomogeneous microstructure on the global component behavior, the mechanical properties of a topology optimized component were tested.Sample ProcessingMaterialAlSi10Mg powder supplied by m4p materials solutions GmbH (Feistritz i.R., Austria) with a powder bulk density of 1.51 g/cm³ served as base material. Table 1 shows the chemical composition of the powder with a size distribution of D10 = 20.5 µm, D50 = 33.9 µm and D90 = 56.8 µm.Table 1: Chemical composition of the AlSi10Mg powder particles in wt.%.Al BaseFe 0.16Si 10.5Mg 0.31Mn Cu <0.01 <0.01 <0.01 <0.01 <0.01 <0.01 <0.01TiZnPbSnPBF-LB systemThe bulk material, which was analyzed in this study, was manufactured with a Creator from O.R. Lasertechnologie GmbH (now: 2oneLab GmbH, Darmstadt, Germany).

Lasertechnologie GmbH (now: 2oneLab GmbH, Darmstadt, Germany). The PBF-LB system is equipped with an Ytterbium fiber laser with a wavelength of 1070 nm and a maximum power of 250 W. The round base plate has a diameter of 110 mm. An F-Theta lens ensures a constant spot diameter across the processing area. During the building process, an Argon atmosphere with a remaining oxygen content of <0.1% in the processing chamber prevents oxidation. Optimized parameters from a previous study with a laser power of 250 W, a scanning speed of 900 mm/s, a hatch spacing of 150 µm, a layer thickness of 30 µm and a spot diameter of 40 µm were used for the PLF-LB process. The buildup process took place without preheating of the base plate. A common bidirectional line scanning strategy with a triple contour scan and a rotational angle of 67° of the hatching pattern of two subsequent layers was used.Sample geometryTwo groups of samples were evaluated in this study, which are summarized in Table 2.

To study geometric influences on the microstructure and the mechanical behavior, the first group consisted of representative geometric elements. The rectangular samples with a height of 50 mm were built with different cross sectional areas of 6x1, 6x2 and 6x6 mm² under varying construction angles of 45, 60 and 75° without the use of support structures. In the second group of samples, three-point bending beams with an optimized topology were manufactured as a demonstrator to study the global component behavior. The two dimensional design proposal for the bending beam was generated using the solver of the software Tosca 2019 from Dassault Systems (Vélizy-Villacoublay, France).

The two dimensional design proposal for the bending beam was generated using the solver of the software Tosca 2019 from Dassault Systems (Vélizy-Villacoublay, France). An area reduction of 40% in a design space of 48x12 mm² was set as a boundary condition with the aim of a maximum stiffness of the part [9].449NiTable 2: Overview of the samples studied.Group A – Bulk materialConstruction angles45, 60 ,75°Cross sectional areas6x1, 6x2, 6x6 mm²Experimental MethodsMicrostructural analysis Small-scale tensile testsGroup B – Bending beams Dimension48x12x12 mm³ PBF-LB PBF-LB+EDM Micro-computed tomography Component testsManufacturing routeExperimental MethodsFor a stable position of the bending beams during the component test, the design proposal obtained from the topology optimization was provided with shoulders at the sides.

The derived design proposals were extracted to a constant thickness of 12 mm, which enables the manufacturing with two different routes, as depicted in Figure 1.Figure 1: Manufacturing of the bending beams analyzed in this study.To ensure manufacturability over both routes, a minimum wall thickness of 1.25 mm was additionally set as a boundary condition to the topology optimization algorithm. In route A, the bending beams were directly built in the PBF-LB process. Since the geometry of the design proposal contains regions which are built under different construction angles, as well as areas of different cross sectional areas, an inhomogeneous microstructure is expected within the parts. To avoid distortion, support structures were used on downskin surfaces with a construction angle of less than 45°. In route B, first a rectangular block of bulk material was manufactured in the PBF-LB process. The contour of the beam was then generated using electric discharge machining (EDM).

In route B, first a rectangular block of bulk material was manufactured in the PBF-LB process. The contour of the beam was then generated using electric discharge machining (EDM). Therefore, a homogeneous microstructure is expected within the parts.Experimental MethodsMicrostructural analysisFor microstructural analysis, representative samples of group A were ground with SiC paper and then polished using diamond pastes. Etching was performed in an aqueous solution of 2% NaOH for 85s. For evaluation of the microstructure, scanning electron microscopy was used. All of the images were captured normal to the building direction. In order to be able to resolve spatial inhomogeneities, the microstructure was analyzed in varying positions within the cross sections of the samples.

In order to be able to resolve spatial inhomogeneities, the microstructure was analyzed in varying positions within the cross sections of the samples. Since gradients in microstructure can also occur across a single melt pool [10], the images were captured in the melt pool center to ensure comparability.Tensile testingThe effects of inhomogeneous microstructure on mechanical material behavior were evaluated using tensile tests. The geometry of the tensile bars is shown in Figure 2. The tensile bars were taken at defined distances from the downskin surfaces of the printed rectangular samples in450group A using EDM. The tensile tests were performed at room temperature using an universal testing machine Z2.5 TN from ZwickRoell GmbH (Ulm, Germany) with a maximum load of 2.5 kN. The tests were conducted at a constant strain rate of 10-3/s until fracture. A preload of 10N was applied.

The tests were conducted at a constant strain rate of 10-3/s until fracture. A preload of 10N was applied. The strain was captured with digital image correlation using the software GOM Aramis (Carl Zeiss GOM Metrology GmbH, Braunschweig, Germany).Figure 2: Geometry of the small-scale tensile bars in mm (left) and positions of the tensile bars in the printed bulk material (right).Component behaviorIn order to investigate the influence of inhomogeneous microstructure on the component behavior, three point bending tests were performed on the samples of group B (see Figure 3). The bending tests were carried out on an universal testing machine from ZwickRoell with a maximum load of 100 kN. The distance between the supporting pins was 48 mm. The loading pins and the supporting pins were rigid and had a diameter of 10 mm. The displacement of the specimens was determined with a tactile displacement transducer, which was placed centrally under the bending beams.

The displacement of the specimens was determined with a tactile displacement transducer, which was placed centrally under the bending beams. The tests were performed at a displacement speed of 1 mm/min.Figure 3: Experimental setup for the performed component tests.Micro-computed tomographyIn order to investigate the influence of defects on the component behavior, the fabricated bending beams were analyzed with micro-computed tomography (µCT). Therefore, the μCT-System PRECISION, developed by YXLON International GmbH (Hamburg, Germany) was used. The X-ray beam was generated at an acceleration voltage of 165 kV and a target current of 0.06 mA. The intensity of the X-rays was recorded using a Perkin Elmer XRD1620 AN flat panel451detector featuring a pixel pitch of 200 µm and an overall size of 2048 x 2048 pixels. The scanning was performed with a focus object distance (FOD) of 51.65 mm and a focus detector distance (FDD) of 913.92 mm, resulting in a voxel size of 11.3 µm.

The scanning was performed with a focus object distance (FOD) of 51.65 mm and a focus detector distance (FDD) of 913.92 mm, resulting in a voxel size of 11.3 µm. In each scan, 1860 projections were captured. The reconstruction of 3D images was performed through a filtered back projection (FBP) algorithm implemented in VGStudioMAX from Volume Graphics International GmbH (Heidelberg, Germany). For qualitative assessment of defect hotspots in the samples, projections of the minimum gray values were generated along the building direction using the thick slab algorithm of VGStudioMAX.Results and discussionMicrostructural analysisIn Figure 4, two micrographs of a bulk sample with a dimension of 6x2 mm² and a construction angle of 45° are presented.

The micrograph at the left was taken at 200 µm to the downskin surface, the one at the right at 1000 µm.Figure 4: Gradient in Microstructure due to a local change of the cooling conditions.Both micrographs show a cellular structure of primary crystallized Al and eutectic Si, which is typical for PBF-LB AlSi10Mg in the as-built state [2, 3]. However, the cellular structure in the left micrograph appears much coarser. It is commonly known from literature, that the cellular size in Al-Si alloys is strongly correlated to the cooling conditions, which can be expressed using the following formula [11]:−𝑛𝑛𝜆𝜆 = 𝑎𝑎(𝐺𝐺 × 𝑅𝑅)is the cooling rate during The product of the thermal gradient and the growth rate solidification, are positive constants. Therefore, it can be concluded that areas with a bigger cellular size solidified at a lower cooling rate. As it is depicted in Figure 4, the heat dissipation during PBF-LB is limited especially on the downskin surface of overhanging parts.

As it is depicted in Figure 4, the heat dissipation during PBF-LB is limited especially on the downskin surface of overhanging parts. Therefore, a heat buildup is formed, which finally results in a local reduction of the cooling rate and thus a coarser microstructure [6, 7]. To analyze how strongly the effect affects different component geometries, the cell size of the bulk samples of group A was evaluated. For this purpose, microstructural images were taken from defined distances to the downskin surfaces. To ensure comparability, the images were taken from the melt pool center since the microstructure can change over the cross section of a melt pool [10]. To evaluate the cellular size, the captured microstructural images were first binarized with ImageJ. The evaluation of the cellular size then was performed with the ImageJ plugin MorpholibJ [12].

To evaluate the cellular size, the captured microstructural images were first binarized with ImageJ. The evaluation of the cellular size then was performed with the ImageJ plugin MorpholibJ [12]. Figure 5 shows the results for specimens at different construction angles (left) as well as for different cross-sectional areas (right).is the cellular size,and(𝐺𝐺 × 𝑅𝑅)𝜆𝜆𝑎𝑎𝑛𝑛452Figure 5: Influence of the construction angle (left) and of the cross sectional areas (right) on the cellular size in PBF-LB AlSi10Mg.All samples show an increasing cellular size towards the downskin surface.

This effect is amplified for a reduced construction angle as well as for a reduced wall thickness due to the limited heat dissipation during the PBF-LB process.Figure 6: Representative stress-strain curves (left) and derived yield strength, tensile strength and elongation after fracture (right).453Tensile testingIn Figure 6 (left), representative stress-strain curves of tensile bars taken from defined distances from the downskin surface of a 6x2 mm² block built up at 45° are depicted. Figure 6 (right) shows the yield strength, tensile strength and elongation after fracture. The results the standard were derived out of DIN EN ISO 6892-1 [13]. The stress-strain curves show an increase in yield strength and tensile strength with increasing distance from the downskin surface. This behavior can be related to the graded microstructure shown in the microstructure analysis.

This behavior can be related to the graded microstructure shown in the microstructure analysis. The cellular structure serves as barrier for dislocations and therefore contributes directly to the yield strength, which can be estimated with the Hall-Petch relationship [14, 15]. Furthermore, the inhomogeneous cellular size indicates changing solidification conditions across a component. Therefore, other strengthening mechanisms, such as the formation of Si precipitates and supersaturation in the solid solution can be influenced as well [14]. Figure 7 shows the influence of the construction angle (left) and the influence of the exposed cross-sectional area (right) on the local yield strength. The results show clearly that a smaller construction angle and a smaller exposed cross-sectional area lead to a larger gradient of the yield strength within one part.

The results show clearly that a smaller construction angle and a smaller exposed cross-sectional area lead to a larger gradient of the yield strength within one part. In general, the determined values for yield strength and tensile strength are comparable with the results of macroscopic tensile tests of PBF-LB AlSi10Mg from the literature [16]. However, the elongation after fracture does not show a clear correlation, as it is shown in Figure 6. It can be assumed that for Al-Si-Mg alloys, defects have a major influence on the elongation after fracture [17]. This is also shown on a representative optical image of the fracture surface of a tensile specimen (see Figure 8).

This is also shown on a representative optical image of the fracture surface of a tensile specimen (see Figure 8). Since the gauge volume of the tensile specimens is very small, variations in porosity between individual samples are to be expected, which therefore explains the higher variations regarding the elongation after fracture.therecorded stress-strain curves accordingtoFigure 7: Influence of the construction angle (left) and of the cross sectional areas (right) on the local yield strength in PBF-LB AlSi10Mg.454Figure 8: Optical image of the fracture surface of a small-scale tensile specimen with marked pore.Micro-computed tomographyFigure 9 shows representative µCT images for the bending beams of each production route. To highlight the porosity, these are projections of the lowest gray values over 5 mm, which were generated using the thick slab algorithm implemented in VGStudioMAX.

To highlight the porosity, these are projections of the lowest gray values over 5 mm, which were generated using the thick slab algorithm implemented in VGStudioMAX. For the bending beam, which was manufactured directly in the PBF-LB process, a global porosity of 0.16% was determined. The PBF-LB + EDM bending beam shows a higher porosity of 0.62%. Due to the exposure of the larger surface area in the PBF-LB process, more process by-products such as spatter are formed. These can be larger than the powder particles and therefore lead to the formation of lack-of-fusion defects [18, 19]. Accordingly, the top and bottom surfaces of the beam produced directly in the PBF-LB process show locally increased porosity.

Accordingly, the top and bottom surfaces of the beam produced directly in the PBF-LB process show locally increased porosity. The thin-walled component areas, on the other hand, exhibit low porosity regardless of the build angle.Figure 9: Projections of the lowest grey values over 5 mm from µCT images of representative bending beams.455Component behaviorThe effects of the geometry-dependent mechanical material behavior were tested by means of component tests on three-point bending beams. Figure 10 shows the force-displacement curves of the tested components, Table 3 summarizes the results.1614N kn1210ie c r o F8642PBF-LB PBF-LB + EDM0012345Displacement in mmFigure 10: Force-displacement curves of the bending beams tested in this study.Table 3: Results of the component tests.Manufacturing route Max. load in kNPBF-LB PBF-LB + EDM13,98 ± 0,26 14,74 ± 0,20Max.

load in kNPBF-LB PBF-LB + EDM13,98 ± 0,26 14,74 ± 0,20Max. displacement in mm 2,03 ± 0,13 2,17 ± 0,10Since the use of different routes to manufacture the beams can lead to deviations in terms of dimensional accuracy and thus influence the deformation behavior, comparability between the two manufacturing routes must be ensured. A corresponding check of the masses revealed a negligible difference of less than 0.5% between the bending beams of both manufacturing routes. Furthermore, the geometries of the bending beams were evaluated on the basis of the obtained µCT data. A surface analysis revealed maximum deviation of 0.1 mm between the two geometries for the upskin and vertical surfaces. Just at downskin surfaces, some local deviations of up to 0.3 mm were detected which could be attributed to powder adhesions on the beam, which was manufactured directly using PBF-LB.

Just at downskin surfaces, some local deviations of up to 0.3 mm were detected which could be attributed to powder adhesions on the beam, which was manufactured directly using PBF-LB. The results obtained by the component tests, show clearly that the PBF-LB + EDM type bending beams have a higher strength than the bending beams built directly in the PBF-LB process, the maximum displacement appears to be similar. This behavior is not in accordance with the findings of the µCT analysis, therefore the microstructure needs to be taken into account. The design of the considered bending beams resulting from the topology optimization contains many thin-walled struts as well as many overhanging component areas. From the investigations carried out within the scope of this work, it can be concluded that different temperature fields are present in the structure of these component areas in the PBF-LB process.

From the investigations carried out within the scope of this work, it can be concluded that different temperature fields are present in the structure of these component areas in the PBF-LB process. Particularly at the downskin surfaces456of thin-walled component areas, heat buildup occurs due to the insulating effect of the surrounding metal powder. The associated locally reduced cooling rate results in a coarsening of the microstructure, which ultimately influences the mechanical behavior. In the case of the bending beams produced via the PBF-LB+EDM route, the generation of the component contour took place after the PBF-LB process, so that no overhanging or thin-walled areas were built up in this route.ConclusionsIn this study, the influence of inhomogeneous material properties on the mechanical behavior of additively manufactured components made of AlSi10Mg was investigated.

For this purpose, representative geometric elements with different wall thicknesses were manufactured under different construction angles at first. A microstructural analysis revealed a coarsening of the microstructure towards the downskin area, which could be associated with a reduced dissipation of the process heat. By tensile tests on small-scale tensile specimens, the mechanical material behavior could be spatially resolved and related to the results of the microstructure analysis. In the second part of the study, the global component behavior was considered.

Materials (Basel, Switzerland) 11, 12.[6] Liu, M., Takata, N., Suzuki, A., and Kobashi, M. 2020. Development of gradientmicrostructure in the lattice structure of AlSi10Mg alloy fabricated by selective laser melting. Journal of Materials Science & Technology 36, 106–117.457[7] Yue, D., Li, D., Zhang, X., Chen, B., Qin, R., and Wang, Z. 2023. Tensile behavior ofAlSi10Mg overhanging structures fabricated by laser powder bed fusion: Effects of support structures and build orientation. Materials Science and Engineering: A 880, 2, 145375. [8] Englert, L., Czink, S., Dietrich, S., and Schulze, V. 2022. How defects depend on geometryand scanning strategy in additively manufactured AlSi10Mg. Journal of Materials Processing Technology 299, 117331.[9] Holoch, J., Lenhardt, S., Revfi, S., and Albers, A. 2022.

perimental Research on Fabrication of Iron Based Alloy and Nano-Al2O3 powder parts by Laser Sintering Jie liu School of Mechanical and Electrical Engineering, Shenzhen Polytechnic Shenzhen, Guangdong, 518055, CHINA AbstractExperiments on selective laser sintering of Iron Based Alloy and nano-Al2O3 Ceramic Bulk Materials are carried out and effect of sintering parameters on t he process is analyzed systematically. A reasonable selective laser sintering technique which can be used to fabricate parts wit h free shape is obtained and verified wit h a multilayer sintering experiment. The component and t he microstructure of t he sintering production is tested. The influences of parameters and the amount of nano-Al2O3 on microstructure and microhardness of the sintering parts are studied. Laser sintering iron-based alloy experiments show that: microhardness has been noticeably improved.

Laser sintering iron-based alloy experiments show that: microhardness has been noticeably improved. It is indicated that with the selective laser sintering technique obtained above, nano-alumina can be processed to manufacture three-dimension parts with free shape. With the addition of Al2O3 and the increase of composite parts of the grain gradually thinning, microhardness gradually improved nanocomposite parts for the microstructure of the dendrite skeleton-shaped crystal and the plane together, the internal Al2O3 dispersion organizations to strengthen the implicit crystal martensite and ferrite mixed organizations.IntroductionRapid Prototyping (RP) can be defined as a group of techniques used to quickly fabricate a scale model of a part or assembly using three-dimensional computer aided design (CAD) data. Rapid Prototyping has also been referred to as solid free-form manufacturing; computer automated manufacturing, and layered manufacturing.

Rapid Prototyping has also been referred to as solid free-form manufacturing; computer automated manufacturing, and layered manufacturing. RP has obvious use as a vehicle for visualization. In addition, RP models can be used for testing, such as when an airfoil shape is put into a wind tunnel. RP models can be used to create male models for tooling, such as silicone rubber molds and investment casts. In some cases, the RP part can be the final part, but typically the RP material is not strong or accurate enough. However, most of the commercially available materials such as photo-polymers, powders, paper, wax, plastic materials, and even rubber, are only for the application of concept models, visual prototypes, and limited functional prototypes. The physical properties of these materials are not suitable for functional prototypes or tooling applications[1-3].

The physical properties of these materials are not suitable for functional prototypes or tooling applications[1-3]. Amongst these materials, only powder sintering has great potential applicability for functional prototypes and it even has potential applicability for direct tooling fabrication. Hence, several researchers have concentrated on the development of the direct laser sintering of metals [4-,6], alumina with polymer binders [7], and metal with polymer binders or with low- melting point components[8,9]. LENS (laser-engineered net-shaping) is the only current commercial direct metal fabrication system .Jet technology is another important process for RP systems. By using a nozzle to squeeze low-melting point alloys or a powder binder mixture, metallic or ceramic parts can possibly be fabricated for the application of functional prototypes or tooling.The objective of laser cladding is to fuse an alloy onto the surface of a substrate with minimum dilution of the substrate.

Areas are usually clad by overlapping single clad tracks [12]. Laser cladding was originally developed for surface treatment to modify surface wear or corrosion properties. A new material was clad on the substrate with metallurgical bonding and low dilution [13]. Usually, only one layer was fused on the substrate in the conventional laser cladding processes. In fact, laser cladding can be considered as one type of material ingressmanufacturing technology, if multiple layers and any designed pattern cladding are possible. A theoretical investigation of the influence of the process parameters on laser cladding's geometrical properties was conducted using thermal modeling and computer simulation.

A theoretical investigation of the influence of the process parameters on laser cladding's geometrical properties was conducted using thermal modeling and computer simulation. High power laser diodes were coupled with fiber optics to fabricate solid freeform parts from metal powder.For the hybrid processes presented in this paper, metal powder was directly delivered into the laser generated melted pool rather than sintered on the powder bed, and the surface or the edge of the clad mold is further machined to achieve the expected accuracy by milling. Hence, the layer manufacturing technology is employed to build up the material of the part. 1. Experimental ProcedureThe design and construction of the hybrid processes of SLC and milling is shown in Fig. 1.Fig. 1 CO2 laser cladding schematic drawingA 3000W CO2 laser was employed in the SLC system as the power resource for the laser cladding process.

1.Fig. 1 CO2 laser cladding schematic drawingA 3000W CO2 laser was employed in the SLC system as the power resource for the laser cladding process. A four axial PC-based controller was used to position the work piece XY table, laser focused lens, and milling head. The focused lens and milling head were moved up layer-by- layer.Fig 2 Layer claddingFig 3 Scan Path of SLCNano-Al2O3 and iron powder were used in this test. Beside of the diamond, Al2O3 has a highest hardness. Nano-Al2O3 reinforcement ceramic or rubber is several times higher than conventionalAl2O3 particles, in particularly, to improve the density of ceramics, finish, cold and heat fatigue, fracture toughness is particularly significant. This study used the nano-particle size for alumina 50 ~ 180 nm, purity ≥ 99.99%, specific surface area of 14±5m2 /g, phase α, α- Al2O3 has some special properties, such as stability, high hardness, stability size.

Fig.6 shows the morphology of cladding layer in different Al2O3 content. The equiaxed grains with 0.5% Al2O3 (in fig. 3a) are coarse. The grains size of cladding decrease with the increasing of the content of Al2O3. Figure 6a is the equiaxed, Figure6b is the dendrite, the microstructure of the sample in the erosion of corrosion is very clearly, and Figure 6c is the combination of dendrite and cell-like crystals.50μm50μm（a）（b）50μm（c） Fig.6 the morphology of indifferent Al2O3 content （a）0.5% Al2O3 （b）1.0% Al2O3（c）1.5%Al2O3Fig 7 is the SEM photo of nano composite coating of 1.5 percent Al2O3; The experimental results show that the grain size is very fine: the primary dendrite spacing is 3 um, the secondary dendrite arm spacing is 6 um. the grain size of alloy is apparently refined. Nano-Al2O3 distributed in the dendrite arm stem and dendrites, while nano-Al2O3 in part has certain level aggregate phenomenon.Fig.7 SEM photoFig.8 shows the micro hardness of part.

PERIMENTAL STUDIES IN STEREOLITHOGRAPHY RESOLUTIONBenay Sager, David W. Rosen, Meghan Shilling and Thomas R. Kurfess The Woodruff School of Mechanical Engineering Georgia Institute of Technology Atlanta, GA 30332-0405ABSTRACTAs we move towards micron-scale rapid manufacturing, it is critical to understand build resolution of Stereolithography technology. In order to determine the resolution limitations, positive and negative features on Stereolithography parts were built and analyzed. Results from several experiments were compared to an analytical model and important resolution issues are highlighted. Based on these experimental results, parameters that will maximize build resolution for a number of well-understood shapes are suggested in the paper. Build resolution experimental results, analysis, and measurement techniques are discussed. Conclusions are drawn related to feature shape as resolution limits are approached.1.

Build resolution experimental results, analysis, and measurement techniques are discussed. Conclusions are drawn related to feature shape as resolution limits are approached.1. INTRODUCTIONStereolithography (SLA) is a layered rapid prototyping process in which an UltraViolet (UV) laser is used to selectively cure a vat of liquid photopolymer resin in order to physically make a part from the generated CAD model. Even though the technology has been commercially available for over 15 years, no real attempt has been made to apply it to micro manufacturing until recently. With the interest in applying this technology to the micro manufacturing area comes the need to study the resolution of Stereolithography. More specifically, the limits of the resolution, both theoretical and empirical, need to be established.It is important to make the distinction between resolution and accuracy before moving any further.

Accuracy is defined as the measure of closeness to the nominal dimensions and geometry of the intended SLA part (Davis 2001). On the other hand, resolution refers to the fineness of the technology, quantifying the absolute limit of the technology.2. STEREOLITHOGRAPHY RESOLUTIONStereolithography resolution is affected by pre-build and build software (Software Imposed Parameters or SIP) and parameters that are inherent in the process (Process Parameters or PP). More information about the classification of resolution-affecting parameters could be found in (Sager and Rosen 2002). SLA resolution can be separated into horizontal (x-y plane) and vertical (z plane) resolution issues, and this paper will focus on the effect of horizontal resolution issues.

SLA resolution can be separated into horizontal (x-y plane) and vertical (z plane) resolution issues, and this paper will focus on the effect of horizontal resolution issues. The resolution limiting factors in horizontal domain are: · · Resolution of the·Linewidth compensation (SIP) Laser beam spot diameter (PP)Stereolithography grid (PP).stl file (SIP)The laser beam spot diameter and the linewidth compensation parameter are the two parameters that have the most influence on horizontal resolution of the process on mesoscale manufacturing. Simply put, the laser beam spot diameter on the build surface is the smallest feasible feature size that can be scanned. Linewidth compensation determines how the laser beam should be offset when scanning a feature so that a cross-section would not be too small or too large. A brief description of these two parameters will be given.

A brief description of these two parameters will be given. However, it is important to keep in mind that all four parameters would have significant influence when micron-size manufacturing is concerned.702.1 Laser Beam Spot DiameterThe laser beam diameter is fixed at 200 (cid:80)m on the build surface for most 3D Systems’ SLA systems. In theory, this is the absolute limiting size of a feature that can be built. However, because of the scanning patterns, the smallest feature that can be built is around one-and-a-half times the laser beam diameter, around 300 (cid:80)m. In high resolution systems such as the 3D Systems’ Viper Si2 machine, the laser beam spot diameter is 75 ± 15 (cid:80)m, and features as thin as 80 (cid:80)m have been built using this spot size.2.2 Linewidth CompensationThe linewidth compensation parameter ensures that the cross section of thin parts is scanned correctly by shifting over the center of the laser beam spot.

In other words, linewidth compensation compensates for the width of the cured linewidth.The default value for linewidth compensation is 125 (cid:80)m, which is ½ of the laser beam diameter. When a rib that is 250 (cid:80)m is to be built, if the default value of linewidth compensation is applied, the outline of the rib is shifted towards the inside of the profile and a line is created rather than a closed profile. Since a line is not interpreted by the SLA system as part of a cross- section, it would not be scanned. On the other hand, if a smaller linewidth compensation value such as 25 (cid:80)m is used, the rib feature will be scanned. This would result in a thicker-than-desired rib that is 400 (cid:80)m thick; however, the feature would not be altogether omitted. However, as aforementioned, it could also result in very inaccurate thin features.3. ANALYTICAL BACKGROUNDStereolithography is a complex process that is enabled by a number of mechanisms working together.

ANALYTICAL BACKGROUNDStereolithography is a complex process that is enabled by a number of mechanisms working together. Because of this, it is important to identify the main components that have a bearing on the theoretical resolution for the process.As the laser beam is scanning along the vat surface, a cured line is formed. The shape of thisline is a parabola (Jacobs 1992), and its cure depth is given as:C (cid:32)dDpE ln( max/E c)(Equation 1)where Cd is the cure depth, Dp is the penetration depth for a resin, Emax is the maximum centerline laser exposure on the build surface, and Ec is the critical exposure for a resin. Equation (1) is known as the working curve for SLA, where Dp and Ec are resin constants. In the horizontal platform, the width of the cured line depends on the laser beam spot radius as well as a resin constant:DCWL o p2/(cid:32)wd(Equation 2)where Lw is the cured linewidth and Wo is the laser beam spot radius.

These values are important to keep in mind when the limits of resolution are considered.4. EMPIRICAL STUDYTo quantify the effects of laser beam spot diameter and linewidth compensation on horizontal build resolution, a number of experiments were performed with different part cross-sections. Of particular interest was the predictability of negative and positive features. After building these cross-sections in Stereolithography machines, they were measured using a number of measurement techniques. The description of each cross-section, its outline, what it was intended to test, and how it was measured are shown in Table 2.Table 2.

The description of each cross-section, its outline, what it was intended to test, and how it was measured are shown in Table 2.Table 2. Shapes used for empirical testingShapeRibsGridsGearsHolesOutlineDetail(units in (cid:80)m)What it is testingHow it is measuredResolution for positive featuresMicrometerResolution variation with build location White light InterferometerResolution variation with curved surfaces/featuresOptical microscopeResolution for negative featuresWhite light interferometer72The detail drawings show the horizontal cross-section for each shape as seen from the top view. The ribs were built using a number of different machine/resin combinations in order to quantify the resolution of different machines. The thickness of the ribs varied from 2.5 mm to 0.25 mm with 0.25 mm increments.The rectangular grids were built in order to quantify how the thickness of a wall varies with respect to changing location on the build platform.

The grids were built at the center of the vat and at the corners. The intended wall thickness for each grid was 50 (cid:80)m.A spur gear with 20 teeth was selected as a suitable profile for resolution study; not only because gears are very important in engineering, but also they have curved positive and negative profiles. The spur gear had 6.25 mm pitch diameter with an intended gap of 0.237 mm between teeth and tooth thickness of 0.179 mm.In order to test the building ability of negative features, a part with holes was designed. Thispart had vertical holes that ranged from 1.25 mm to 0.125 mm in diameter with 0.125 mm increments. In all the parts built, the effect of the linewidth compensation parameter was also studied.4.1 Data Collection and AnalysisDepending on the outline of the cross-section measured and the level of measurement detail required, different measurement techniques were used in this study.

For measuring the thickness of the rib parts, using a micrometer was sufficient. For the gear cross-sections, the outline of the gears as well as the measurements were desirable. Therefore, an optical microscope with a 5x zooming lens was sufficient.Interferometers, in particular white light interferometers, are currently being used to inspect mesoscale parts. Georgia Tech has a Zygo NewView 200 white light interferometer, which was used for collecting much of the grid and hole data. The interferometer uses the principle of interference to create a 3-dimensional surface profile of the parts. This data, however, is limited to a single view due to the slope limitations on measurement. Because the white light interferometer was initially created to measure surface characteristics such as surface roughness, it was designed to have very good resolution in the z-direction (out of plane). This resolution is more than adequate for mesoscale parts.

This resolution is more than adequate for mesoscale parts. The resolution in the lateral (x and y) directions, however, is not quite as good but will be sufficient for initial measurements. The lateral resolution is dependent on the objective being used. For the objectives in the Precision Machining Laboratory the lateral resolutions are 2.2 (cid:80)m and 8.8 (cid:80)m (for the 10x and 2.5x objectives, respectively) at regular camera resolution.Conventional coordinate metrology is not appropriate for the type of data that are gathered using a white light interferometer. A single surface does not hold enough information for comparison to a CAD model or to an analytic surface (other than a plane). Because a CAD model is a series of surfaces that intersect, data from other surfaces are needed to properly orient the model or surface to the data.

Because a CAD model is a series of surfaces that intersect, data from other surfaces are needed to properly orient the model or surface to the data. Due to this limitation, two-dimensional curves such as lines and circles are of interest.73For two-dimensional analysis, the data on a single surface are not what need to be analyzed. The edges of the data hold the important information. Therefore, it is necessary to extract edge data points from the original point cloud before any dimensional analysis can be performed. The edge point extraction takes several steps. First, a plane is fit to the surface data and an “image” is created. The pixel size of the image is set to the distance between the evenly spaced data points and the intensity, or color, of the pixel is directly related to the height of the pixel above the fit plane.

The pixel size of the image is set to the distance between the evenly spaced data points and the intensity, or color, of the pixel is directly related to the height of the pixel above the fit plane. After the “image” is created, a median filter is applied to remove noise.After being filtered, edge points are either located by binarization and a simple row and column sweeping algorithm or an edge detector and contour tracking algorithm based on the idea of finding the center of mass in combination with a contour tracking technique. The edge points can then be analyzed by being fit to two-dimensional shapes such as lines, circles, and ellipses.

The edge points can then be analyzed by being fit to two-dimensional shapes such as lines, circles, and ellipses. This entire process is illustrated in Figure 1.Image Image Creation CreationBinarization BinarizationEdge Point Edge Point Location: Location: Sweeping SweepingFiltering FilteringEdge Edge Analysis AnalysisEdge Edge Detection DetectionThresholding ThresholdingEdge Point Edge Point Location: Location: Contour Contour Tracking TrackingFigure 1: Analysis FlowchartThe methods described above have been validated using generated test data of circles and bars. Perfect circles and perfect bars yielded perfect results. When known noise was added, both sets yielded expected results. This procedure is fully outlined in (Shilling 2003).4.2 Rib PartsThe outline of the rib parts and their detail cross-section were presented in Table 2.

This procedure is fully outlined in (Shilling 2003).4.2 Rib PartsThe outline of the rib parts and their detail cross-section were presented in Table 2. The rib parts were built using a number of machine/resin combinations, linewidth compensation, and layer thickness values, which are shown in Table 3. The goal was to quantify the resolution for each machine, along with the effect of different linewidth compensation values. In short, four comparative studies were done with the rib parts:1. Thickness at tip of rib versus thickness at base of rib 2. Thickness of rib on top surface of build versus thickness of rib on bottom surface of build 3. Resolution of Stereolithography machine used 4. Resolution with respect to linewidth compensation value usedThe thickness values at the tip of ribs were very similar to those at the base of ribs. It should be kept in mind that the distance between the tip and base of each rib is very short, about 2.5 mm.

It should be kept in mind that the distance between the tip and base of each rib is very short, about 2.5 mm. Therefore, such a small change in location of the laser beam as it was scanning the parts would not cause a significant change in the resolution of the final parts.It is known that upfacing and downfacing surfaces in SLA have different resolution. This is partly due to the curing effects, and partly due to the fact that at the bottom of each build support74structures are anchored to the build platform (Jacobs 1996). For all six rib parts built, the thickness for a particular rib (rib 5) at the top and bottom surfaces is shown in Figure 2. There is a consistent thickness difference between bottom and top surfaces. This thickness difference is quantified in Table 4.Table 3. Build parameters for rib parts 1-6Part Machine1 2 3 4 5 6SLA 3500 SLA 250 SLA 250 Viper Si2 Viper Si2 Viper Si2ResinSL 7510 DSM 7110 DSM 7110 SL 5510 SL 5510 SL 5510Linewidth Comp.

This is an expected result, since the bottom surface of SLA builds typically have poorer resolution. When a SLA layer is scanned, in order to make sure it would anchor to the previous layer, it is “overcured” by a specified thickness into the previous layer. Additional overcuring of a number of layers causes the bottom surface of a SLA build to have thicker cross-sections than intended; hence a poorer resolution than top build surface. Therefore, it can be concluded that the top surface has features that are on average between 10-20 microns thinner than the bottom surface.Another aspect of this study is quantifying the resolution of different SLA machines used. For each rib, the difference between the intended and measured thickness for different machines75is shown in Figure 3. It can be seen that when the default value for the SLA 250 machine was used, a part whose ribs were thinner than intended was obtained.

It can be seen that when the default value for the SLA 250 machine was used, a part whose ribs were thinner than intended was obtained. On the other hand, as expected, the Viper si2 machine high-resolution build style produced parts with highest resolution.0.30.20.1Intended - Measured (mm)00.10.20.312345678910Rib numberSLA 250SLA 3500SLA Viper si2Figure 3. Average dimensional error for different SLA machinesA comparison of rib width values based on the different linewidth compensation values that can be used when preparing a build for the Viper si2 machine was made. Three different linewidth compensation values were used: the default value of 0.0375 mm, 0.01875 mm, and 0 mm. These different values and their resulting measurements are shown in Table 5.Table 5.

Three different linewidth compensation values were used: the default value of 0.0375 mm, 0.01875 mm, and 0 mm. These different values and their resulting measurements are shown in Table 5.Table 5. Comparison between expected and measured errorLinewidth compensation value used ((cid:80)m) 37.5 18.75 0Analytical expected error ((cid:80)m) 0 37.5 75Average measured error ((cid:80)m) 39 57 104Error discrepancy ((cid:80)m) 39 20 29As shown in Table 5, the lowest average error of 39 microns was obtained when the default linewidth compensation value of 37.5 microns was used. When the same default linewidth compensation value is used, the expected error is zero because the thickness of the laser beam is adequately compensated for. There seems to be a consistent error between the measurements and expected results when different linewidth compensation parameter values are used.

There seems to be a consistent error between the measurements and expected results when different linewidth compensation parameter values are used. It can be argued that by accounting for the error based on linewidth compensation value used, the error between the intended and measured thickness could be reduced. Table 5 presents expected error discrepancy within the range of 20-40 microns based on empirical results. It is clear that the parts built will be thicker than intended, and this value can be quantified as between 20 and 40 (cid:80)m.4.3 Grid PartsThe grid parts were built for a number of studies, including build location, orientation, and effect of linewidth compensation parameter on build resolution (Sager 2003). However, in this paper, only the effect of build location on build resolution will be presented. Four identical test parts were built simultaneously at the four corners of the build platform of the Viper si2 machine.

Four identical test parts were built simultaneously at the four corners of the build platform of the Viper si2 machine. The labeling for the locations of these parts is presented in Figure 4. Each grid part had76the dimensions of 25 mm by 25 mm in x and y directions respectively. The distance between the center of the vat and the center of the grid parts was 112.5 mm.Back of platformFront of platformFigure 4. Location of grid parts on build platformMeasurements of the vertical cross-section of the parts taken by researchers at Siemens AG yielded results that suggest a heavy dependence of wall thickness on build location, as shown in Table 6. These measurements were taken from the grids around the center of each grid part and represent the average thickness of the walls. For each wall, the thickness was measured at three locations. These locations were towards the top of the wall, around the center of the wall, and towards the bottom of the wall.Table 6.

For each wall, the thickness was measured at three locations. These locations were towards the top of the wall, around the center of the wall, and towards the bottom of the wall.Table 6. Grid average wall thickness with respect to build platform locationPosition on platform Center 1 2 3 4Thickness at location ((cid:80)m) 2 1 133 126 265 226 246 210 263 250 217 1713 121 253 274 275 184Representative thickness ((cid:80)m) 127 248 243 263 191Ignoring the smallest and largest values that were obtained at corners 4 and 3 respectively, the representative thickness at the corners can be taken as 243-248 (cid:80)m. When compared to the wall thickness at the center of the build platform of 127 (cid:80)m, this value is almost doubled. Such a difference between wall thickness values at center and corner of the build platform suggests that location on the build platform plays a major role in determining how thick a wall will become.

Such a difference between wall thickness values at center and corner of the build platform suggests that location on the build platform plays a major role in determining how thick a wall will become. This could be attributed to the change in laser beam characteristics as the laser beam moves away from the center of the build platform. For example, in order to adequately calculate the shape of cured line, the changes in the shape and size of the laser beam radius must be taken into account. Moreover, as we move away from the center of the build platform, the laser beam becomes out of focus, causing the wall thickness to become larger.In addition to change in wall thickness with respect to build location, change in wall thickness with respect to vertical location was observed. This phenomenon is a direct result of overcuring each layer and is shown in Figure 5. In Figure 5, the bottom of the wall is the right side.

This phenomenon is a direct result of overcuring each layer and is shown in Figure 5. In Figure 5, the bottom of the wall is the right side. As shown, the thickness at the bottom of the wall is around 130-145 (cid:80)m. The left side of Figure 5, which corresponds to the top of the grid wall, was measured to be around 100 (cid:80)m under the microscope.77Conical TipsFigure 5. Vertical cross-section of grid wallsAs aforementioned, the walls are thinner at the top due to the lack of overcure and print through for the latest-built layers. Print-through is a Stereolithography error that accumulates on a lower layer when an upper layer is scanned. The analytical model expressed by Equation (3) enables an explanation of this cone-tip effect (Rosen 2002; Sager 2003). This model suggests that the wall thickness of the build surface is 78 (cid:80)m and the representative wall thickness is 95 (cid:80)m (Sager 2003) for the Viper si2 high-resolution build style.

This model suggests that the wall thickness of the build surface is 78 (cid:80)m and the representative wall thickness is 95 (cid:80)m (Sager 2003) for the Viper si2 high-resolution build style. This 17 (cid:80)m discrepancy is in the middle of the measured overcure related error range of 10 - 25 (cid:80)m. A representative wall thickness can be computed by adding the analytical thickness estimate to the linewidth compensation error range (20 - 40 (cid:80)m). This yields an expected wall thickness range of 115 - 135 (cid:80)m, which is smaller than the observed thicknesses. Additional research is needed to investigate this difference.Table 7.

This yields an expected wall thickness range of 115 - 135 (cid:80)m, which is smaller than the observed thicknesses. Additional research is needed to investigate this difference.Table 7. Analytical, empirical, and expected empirical values for grid part wallsMeasurement Wall thickness on build surface ((cid:80)m)Analytical Empirical Expected range78 100 78 + (20~40) = 98~118Representative wall thickness ((cid:80)m) 95 130~145 95 + (20~40) = 115~1354.4 HolesA hole study was performed to determine both the resolution of negative features as well as the changes in hole diameter from the top surface of the part to several layers into the part.

In addition to comparing feature sizes in the top and lower layers, the purpose of this test was to determine the smallest hole that could be built.After initial measurement, the part was ground using a small sample grinder to get a surface finish sufficient for measurement by the interferometer, so that the size of holes beneath the top surface can also be measured. From visual inspection, this was determined to be the 0.375 mm- diameter hole. After x-y-z data sets were collected using the interferometer, the data was78analyzed according to the steps outlined earlier in this paper. The results from measurement of the holes, as well as expected results are presented in Table 8.Table 8.

The results from measurement of the holes, as well as expected results are presented in Table 8.Table 8. Radii from Hole Test Parts (in (cid:80)m)PART LOCATIONCircle A (1) B (2) C (3) D (4) E (5) F (6) G (7) H (8) I (9) J (10)Expected (1) 635.00 571.50 508.00 444.50 381.00 317.50 254.00 190.50 127.00 63.50Center (2) 559.67 500.80 569.12 462.41 386.75 319.17 251.93 193.63 122.59 46.09Center ground (3) 614.19 623.16 496.11 446.43 384.09 312.55 247.38 190.50 110.56 -The parts were measured using the Zygo New View 200 white light interferometer with the 10X objective (2.2.(cid:80)m lateral resolution). There were problems with the Zygo stage, so the largest area that could be analyzed was approximately 700 by 500 (cid:80)m. This caused some problems with the circles of larger diameters, but was sufficient for circles of smaller diameters.Not enough data from the largest three holes in the hole test parts was collected to allow for reliable measurement.

Although the results are presented, the values given might not represent the actual values of the radii of the holes. Additionally, the smallest hole was not present after the specimen was ground. The difference between the intended and measured hole radii is presented in Figure 6.201510Intended - Measured ((cid:80)m)5051015203457 Hole Number68910Intended-ungroundIntended-groundFigure 6: Comparison of ground and unground part79The results from comparing the ground part to the same unground part are presented in Figure 6, ignoring the largest three holes because of problems with the interferometer and the smallest hole because of its disappearance. It is seen that the diameters of the holes for the ground part are slightly smaller than that of the unground part in all cases, reflected by the larger difference between intended and measured hole diameters.

This is consistent with the idea that the lower layers are exposed to more UV light; thus have a chance to cure more thoroughly.4.5 GearsA number of 6.25 mm-diameter spur gears were built in order to study resolution of curved cross-sections. The gear had teeth width of 179 (cid:80)m and gap between teeth of 237 (cid:80)m. The average thickness of gear teeth was 186 (cid:80)m on top and 198 (cid:80)m on bottom surface when small feature compensation parameter (SFP) was used. The gap between gear teeth, a negative feature, was 312 (cid:80)m on top and 280 (cid:80)m on bottom surface when SFP was used.It can be said that SFP enabled generation of thin positive features that are more dimensionally accurate on the top surface of a build. In addition, the profile of the gear teeth was much sharper with the use of SFP.

In addition, the profile of the gear teeth was much sharper with the use of SFP. It was noticed on the SLA parts that some of the gear teeth lost their curved involute profile accuracy because they were too thin, which shows the limits of SLA resolution were reached with this spur gear profile.5. CONCLUSIONS AND FUTURE WORKWith the rib parts built, it can be concluded that the bottom surface of SLA parts are thicker than the top surface. For the SLA Viper si2 system, the thickness discrepancy has been measured to be 20-40 (cid:80)m, which is consistent with the analytical model of the curing process. The grid part experiments suggest that as we move away from the center of build platform, the resolution of SLA machines changes dramatically. Based on the empirical results and comparison with analytical models, a range of expected thickness values can be specified for high aspect ratio walls.

Based on the empirical results and comparison with analytical models, a range of expected thickness values can be specified for high aspect ratio walls. Even though the Viper si2 system has a laser beam spot diameter of 75 (cid:80)m, the realistic range for representative wall thickness is 115-135 (cid:80)m. The ability to build small negative features using SLA technology has been demonstrated. However, more work is needed in quantifying the smallest negative feature that can be built. For certain shapes that are well understood, values for certain software parameters that would maximize build resolution are presented in Table 9.Table 9.

For parts that have thin walls or small features, turning the linewidth compensation off will ensure that the small feature is not omitted by the laser. This is critical for all the SLA machines that do not have small feature preservation compensation parameter.The laser beam spot diameter is the biggest contributor to the limitation of SLA resolution. The spot diameter determines the size of the smallest feature that could be built on the horizontal platform. Since SLA is a stacking of a number of two-dimensional layers, it can be said that the resolution in the horizontal direction dictates the resolution in the vertical direction. Therefore, being able to control the laser beam spot diameter would improve the process resolution greatly.For maximum x-y resolution, parts should be built in the center of the build platform if possible. In addition, using the sweeper blade ensures that the resin on the top surface of the SLA build is evenly distributed after each layer is scanned.

In addition, using the sweeper blade ensures that the resin on the top surface of the SLA build is evenly distributed after each layer is scanned. Ribs or any other high aspect ratio features should be built in the vertical orientation to reduce stair steps.Although white light interferometry was originally developed for surface measurements, it can be a useful tool in mesoscale part characterization when combined with several image- processing techniques. This method, however, is quite limited in its lateral resolution, speed, and three-dimensional measuring capabilities. In order to increase the speed and completeness of the measurement, new systems need to be developed.In order to truly quantify the resolution of SLA technology, a study of three-dimensional resolution is needed. In the future, more experiments should be performed with negative features, including shapes that are rectangular such as a square-shaped hole.

In the future, more experiments should be performed with negative features, including shapes that are rectangular such as a square-shaped hole. In addition, a comprehensive cure model that will take into account the effect of focus depth, change of laser beam angle, and resin properties is under study.ACKNOWLEDGEMENTSWe gratefully acknowledge the support from the RPMI member companies and the George W. Woodruff School of Mechanical Engineering at Georgia Tech. This work was partially funded by the National Science Foundation under Grant Number DMI-9988664.REFERENCES Davis, B. E. 2001. "Characterization and Calibration of Stereolithography Products and Processes", Master's Thesis,Georgia Institute of Technology, Atlanta, GA.Jacobs, P. F. (1992). Rapid Prototyping & Manufacturing: Fundamentals of Stereolithography, Society ofManufacturing Engineers.Jacobs, P. F. (1996).

 experimental study of cellular mechanical interface in a bi-material structure fabricated by material extrusion additive manufacturingSumit Paul, Li Yang Department of Industrial Engineering, University of Louisville, Louisville, USAAbstractThis work is a continuation of the preliminary project with the investigation of the characteristics of the cellular structure-based mechanical interlocking interface designs for bi-material structures fabricated by material extrusion additive manufacturing (AM). Three different cellular designs, including auxetic, body centered cubic (BCC), and octahedral, were investigated for the topology design effects on the interface performance. In addition, the effects of build orientation and interface polarity were also studied.

In addition, the effects of build orientation and interface polarity were also studied. The results clearly suggested that the cellular-based mechanical interlocking interface exhibit significantly enhanced ductility and energy absorption that can be desirable to many applications, and that the interface characteristics are influenced by both the topology design of the cellular structures and the intrinsic properties of the processed materials.1. IntroductionMulti-material with additive manufacturing (AM) is being extensively investigated due to the expanded designability with more than one materials available in the structures. Through the introduction of heterogeneous materials, advanced design concepts such as functionally gradient materials (FGM) or metamaterials can not only have locally optimized architectures for performance requirements, but also achieve higher level of design optimality.

On the other hand, the fabrication of multi-material components imposes additional challenge. Various design factors must be considered, such as differences in thermal behaviors of materials, the interfacial integrity, and the interaction of multiple material phases under mechanical and physical loading conditions. Such design framework requires comprehensive consideration of material models, topology designs, as well as the corresponding AM processes, which is currently lacking in general [1].Various AM processes routinely employ multi-material printing, such as material extrusion, material jetting and sheet lamination. Among them, material extrusion is among the most broadly adopted processes, due to its capability to readily print multiple materials via the configuration of multiple printhead nozzles. Various literatures have reported the use of material extrusion for multi-material structure printing in different applications. Watschke et al.

Various literatures have reported the use of material extrusion for multi-material structure printing in different applications. Watschke et al. [2] printed ABS-PLA dual material samples by material extrusion and did tensile testing, lap-shear, compression-shear tests. Yang et al. [3] developed a material extrusion 3D printer for continuous fiber reinforced thermoplastic composites. Khondoker et al. [4] proposed a bi-extruder material extrusion multi- material printer for FGMs made from immiscible thermoplastics and tested mechanically interlocked extrudates of PLA, ABS and polystyrene.2244However, the material option for the multi-material printing with material extrusion is in general limited. Freund et al.

Freund et al. [5] evaluated the peeling resistance and roll peeling resistance of interfaces between a number of common material pairs using material extrusion, including ABS/NinjaFlex, XT/NinjaFlex, XT-CF20/NinjaFlex, GF30-PA6/NinjaFlex, GF30-PP/NinjaFlex, PLA/NinjaFlex, ABS/TPU-93 and PLA/TPU-93. Results showed that with most combinations the interfacial strength is limited. Similar conclusion was also drawn by Lopes et al. [6]. The limitation with material selection results in design limitations. In many cases, material pairings with significantly different properties (e.g. hydrophobic + hydrophilic, soft + hard) are needed for desired functionality, and in these cases the intrinsic material compatibility could become a significant barrier for the adoption of AM technologies.One traditionally-established concept that might provide an effective solution to the multi-material compatibility challenge is the mechanical interlocking design.

The concept of mechanical interlocking design for multi-component structures utilizes the structural interaction between the components to introduce structural integrity. Instead of diffusion bonding at the material interfaces or fastening, the interactions among the components such as contact and sliding result in coordinated deformation, and therefore the mechanical strength of the interlocking design. Mechanical interlocking design has been extensively employed in certain application areas such as woven fabric structure designs. Another perhaps more popular example is the Chinese wooden puzzles. Therefore, the interlocking design mechanism has the potential to overcome the limitation with intrinsic material compatibility and to achieve adequate interfacial strength.There currently exist some literatures that demonstrate the performance of mechanical interlocking multi-material interface designs. Zhao et al.

Zhao et al. [7] investigated mechanical interlocking between A5052 aluminum alloy and 30% glass-fiber-reinforced polybutylene terephthalate joined by injection molding the polymer to the pre-treated textured metal surfaces (Fig.1a). It was found that the strength of the interface is determined by both the injection molding process parameters and the topology design of the surface texture. Rossing et al. [8] investigated an interlocking concept based on grid-like scaffold for PLA-silicone multi-material structures (Fig.1b), and the results showed significant improvement of part integrity despite the low intrinsic material interfacial strength between the two materials. Cellular structure design concepts have also been explored, due to their interconnected porosity that inherently accommodates interlocking design. Perets et al.

Cellular structure design concepts have also been explored, due to their interconnected porosity that inherently accommodates interlocking design. Perets et al. [9] studied the mechanical properties and biocompatibility of a biodegradable implant design with Ti6Al4V cellular structures made by laser powder bed fusion additive manufacturing (L- PBF-AM) and filled with a Mg-based alloy (Fig.1c). Although the study did not focus on the interface design, the design approach resulted in good structural integrity and improved balance of performance. Verma et al. [10] investigated cellular structure-based interlocking design for a 316L stainless steel-AV651 epoxy resin bi-material interface. The cellular structures were designed with several unit cell topologies (BCCZ, CFCC and Kelvin) and manufactured using L-PBF-AM, and subsequently filled with the resin via injection molding (Fig.1d).

The cellular structures were designed with several unit cell topologies (BCCZ, CFCC and Kelvin) and manufactured using L-PBF-AM, and subsequently filled with the resin via injection molding (Fig.1d). The results clearly suggested that the properties of the interface are significantly influenced by the selection of the cellular unit cell topology designs.2245a. Metal-polymer interlocking interface via injection molding over textured surface [7]b. PLA-silicone interface via crossover beam- based interlocking design [3.1]c. Ti6Al4V-Mg multi-material design with cellular scaffold [9]d. 316L Stainless steel-AV651 resin interlocking design with cellular scaffold [10]Fig.1: Mechanical interlocking designs from literatureIn this study, three types of cellular unit cell designs, including octahedral, re-entrant auxetic, and BCC, were investigated for the cellular-based mechanical interlocking designs for the PLA-PETG bi-material interface.

The two materials were chosen due to the expectation of low compatibility as well as material availability at the time of the study. This was also a continuation of the preliminary study that was previously reported [11]. The three types of unit cells were chosen in the attempt to investigate potential significance of representative design rules. The octahedral structure is a stretching-dominated design and exhibits directional isotropy. The re-entrant auxetic structure exhibits negative Poisson’s ratio and is a bending dominated design. The BCC structure is a bending-dominated design and exhibits positive Poisson’s ratios (Fig.2). An experimental study was carried out for the mechanical property evaluation of interfaces based on the three unit cell designs, and the results were also compared to additional finite element analysis (FEA) for insights into the structural behaviors.2. Materials and Methods 2.1.

Materials and Methods 2.1. Design of experimentIn this study, relative density of the cellular design was the only geometry design parameter that was investigated. To minimize the effect of other geometry parameters, the geometry aspect ratio of each of the unit cell designs was arbitrarily decided and kept constant, and the design of relative design was achieved by isotropic scaling of the unit cell dimensions. In addition, the strut thickness of each type of unit cells was also kept constant at all levels of relative densities. Fig.2 shows the unit cell designs, and Table 1 lists the geometry designs of all three types of unit cell structures in2246order to achieve the three levels relative density (0.3, 0.4 and 0.5), as well as the relative density levels.

The relative density levels were determined based on the reference material strength of PLA and PETg from literature, which were set to be about 18MPa and 12MPa, respectively [6, 7].L1L2XYZL3XYZHθ L2L1L2XYZL3a. Octahedralb. Re-entrant auxetic Fig.2 Unit cell design for the cellular structuresc.

Octahedralb. Re-entrant auxetic Fig.2 Unit cell design for the cellular structuresc. BCCUnit cellRe-entrant auxeticBCCOctahedralRelative density (Design/Actual) 0.4/0.397 0.5/0.509 0.6/0.608 0.4/0.395 0.5/0.505 0.6/0.602 0.4/0.399 0.5/0.506 0.6/0.560L1/H (mm) L2 (mm)3.9 3.3 2.9 7.63 6.48 5.72 4.3 3.6 3.32.97 2.51 2.2 7.63 6.48 5.72 4.3 3.6 3.3L3 (mm)6 5.1 4.5 4.3 3.6 3.3θ (deg)75 75 75t (mm)1.3 1.3 1.3 1.9 1.9 1.9 1.3 1.3 1.3Table 1: Design parameters of the cellular lattice structures (positive component)For the interface design, the cellular structures constructed by the unit cell designs (Fig.2) were conveniently termed “positive” component, which were then filled with their complementary geometries that were termed “negative” component, as shown in Fig.3.

The interfaces were designed into the shape of tensile testing coupons, albeit not following the dimension guidance of ASTM E8.Fig.3 Design of the interface samplesIn order to establish comprehensive design of experiment (DOE) space, designs with material assignment of both PLA and PETg to the positive component were included in the investigation. Consequently, the naming of the sample designs follow the custom of “unit cell design”-“relative density (RD) level”-“positive material”, i.e. Aux-40RD-PETg means that the baseline cellular design is an re-entrant auxetic with a 0.4 relative density, and the material for the positive2247component is PETg. Following such DOE, a total of 18 design combinations were included in the study.

Following such DOE, a total of 18 design combinations were included in the study. Due to the resource limitation, 2-3 samples were fabricated for each design.In addition, in order to establish the baseline mechanical properties for the intrinsic material interface, both single-material tensile coupons and bi-material planar interface samples were designs. As shown in Fig.4, bi-material tensile coupon models with the two components overlapping (shown as d in Fig.4) by 0mm, 0.1mm, 0.5mm, 0.7mm and 1mm were designed and fabricated.Fig.4 Benchmark interfacial strength evaluation sample design2.2.Sample fabricationThe samples were fabricated using a FlashForge Creator Pro 3 material extrusion printer. A preliminary process development study was carried out before the fabrication of the actual samples, in order to ensure adequate material affinity that avoids component detachment during the printing.

Such process was done in an ad-hoc way without the concern of achieving truly optimized process settings, therefore is not reported in this paper. Afterwards, the parameters were fixed for all the samples. Table 2 lists the process parameter settings used for printing the samples, and Fig.5 illustrates the fabricated samples (Re-entrant auxetic with 0.4 relative density). All the cellular interface design samples as well as the single-material tensile coupons were fabricated successfully. On the other hand, the benchmark bi-material tensile coupons with 0mm and 0.1mm interface overlap completely failed to maintain part integrity.

On the other hand, the benchmark bi-material tensile coupons with 0mm and 0.1mm interface overlap completely failed to maintain part integrity. The rest of the benchmark coupons were fabricated successfully.Process parameterLayer height (mm)Substrate temperature (°C)Base printing speed (mm/s) # of shells (shell layers) @all sides Infill pattern Infill cross angle (°) Use of brimSetting0.30mm70602Line 90 NoneProcess parameter PLA nozzle temperature (°C) PETG nozzle temperature (°C) Travel speed (mm/s) Maximum perimeter path overlap (%) Infill (%) Infill start angle (°)Setting19522510030100 45Table 2: Process parameter values2248a. Aux-40RD-PLA: PLA (black) in positive componentb. Aux-40RD-PETg: PETg (grey) in positive componentFig.5 Fabricated samples: Re-entrant auxetic with 0.4 relative densityThe samples were subjected to standard tensile testing using an Instron 5569A universal tensile tester with a 5KN load cell.

The displacement of the crosshead was utilized for the calculation of elastic modulus, which was expected to introduce some error. However, due to the low elastic modulus with the structures, this error was not considered significant enough to create biased observations and therefore not further addressed.3. Results and Discussion Fig.6-7 shows the tensile testing results for the single-material coupons. PLA exhibits higher strength and elastic modulus but also more brittle behaviors compared to PETg. From Fig.6, it could also be seen that the large elongation to failure of the PETg component (Fig.7) is primarily attributed to the stretching of the fibrils.

From Fig.6, it could also be seen that the large elongation to failure of the PETg component (Fig.7) is primarily attributed to the stretching of the fibrils. Considering the fact that with material extrusion process the fibrils are defined essentially by the individually deposited tracks, it could also be hypothesized that the process setting for PETg might not be well-tuned in this study, which resulted in relatively low inter-track bonding and consequently failure mode dominated by inter-track delamination. As it is not within the interest of this study to investigate the failure mode of material extrusion constructions, this was not further investigated. Instead, the baseline material properties were utilized for subsequent experimental discussions (Table 3).a. PLAb. PETgFig.6 Single-material tensile coupons after testing2249Fig.7 Stress-strain curves of PLA and PETg couponsMaterialPLA PETgYield strength (MPa) Max.

PLAb. PETgFig.6 Single-material tensile coupons after testing2249Fig.7 Stress-strain curves of PLA and PETg couponsMaterialPLA PETgYield strength (MPa) Max. strength (MPa) Elastic modulus (MPa) 48.76509 36.0392633617.65513 11.80584333823.9766667 553.6733333 Table 3: Mechanical properties baselines of PLA and PETgDuctility (%) 17.65513 35.89736667The tensile testing results for the benchmark interface samples are presented in Fig.8, and a representative sample after failure is illustrated in Fig.9. All the benchmark interface samples failed along the PLA-PETg interface, demonstrating low intrinsic material compatibility. The artificially introduced overlap in the model did facilitate more significant bonding with the interface, although this effect does not appear to scale with the overlapping size.

The artificially introduced overlap in the model did facilitate more significant bonding with the interface, although this effect does not appear to scale with the overlapping size. In addition, although not inspected, it could be expected that the overlap design would result in loss of dimensional accuracy and possibly feature resolution for small-size features for applications.a. Stress-strain curveb. Ultimate strengthc. Elongation to breakaged. Elastic modulus2250Fig.8 Mechanical properties of the benchmark material interfaceFig 9: All benchmark interface samples fractured along PLA-PETG interface (showing 0.7mm-overlap interface)For the tensile testing of the cellular interface samples, the failure is located on the PETg side of the end of the cellular interface section for all the samples, as illustrated in Fig.10 for the Aux- 60RD-PETg sample.

Fig.11-13 shows the stress-strain curves of the cellular interface samples with PLA for the positive component and re-entrant axuetic, octahedral and BCC designs, respectively. It is noted that due to the non-standard gripping end geometry designs and the occurrence of defects with the gripping ends that could be found with some samples, the initial sections of the stress- strain curves for many of the interface samples do not reflect the tension of the sample but instead the slacking of the setup. Therefore, the curves were treated with extra care during analysis of mechanical properties.For the re-entrant auxetic-based design, the behaviors of individual samples exhibit some non- consistency, which might be caused by both process defects and the unit cell topology design. The re-entrant auxetic unit cell contains vertical elements of relatively small thickness (1.3mm), which could potentially cause reduced process quality and large property variability.

The re-entrant auxetic unit cell contains vertical elements of relatively small thickness (1.3mm), which could potentially cause reduced process quality and large property variability. There exists a significant change of mechanical strength with the increase of cellular design relative density from 0.4 to 0.5, and comparatively less change between relative densities 0.5 to 0.6. Further inspection with the fracture surface reveals that this was likely attributed to the change of the fracture mode. For the Aux-40RD-PLA samples, the fracture surfaces exhibit relative smooth and planar tomography (Fig.15), indicating relatively comparative significances of both the PLA and the PETg phases during fracture failure.

On the other hand, for the Aux-50-RD-PLA and Aux-60-RD- PLA samples, fractured tracks that are partially pulled loose from the main structure are clearly present on the fracture surfaces, and the overall surface tomography exhibits more obvious topology-influenced pattern (Fig.14 and Fig 16). This might be due to the reduced volume (are consequently areal) percentage of the relatively weaker PETg phase at 0.5 and 0.6 relative densities, which resulted in PETg-phase-dominated failure with the samples.In comparison, the stress-strain curves of the octahedral-based and BCC-based cellular interfaces are relatively consistent. The octahedral designs exhibit a somewhat more “brittle” behavior, with residual strength of the interface rapidly decreasing with additional strain.

The octahedral designs exhibit a somewhat more “brittle” behavior, with residual strength of the interface rapidly decreasing with additional strain. On the other hand, the BCC designs exhibit higher strength and higher overall ductility, which can be clearly discerned from the long tails of the stress-strain curves (Fig.13) that indicate high residual strength after initial failure. The failure characteristics of both types of interfaces exhibit significant influence from the weaker PETg phase. The fracture surfaces exhibit both PETg phase pull-out and loose PETg tracks (Fig.14 and Fig.16), which indicate that not only does the PETg negative components2251are more significantly loaded, but also that the inter-track bonding of the PETg phase becomes a significant cause of failure.The effects of relative density of cellular structure on the interface property are significant for both types of cellular interface designs, although the trends are different.

For the octahedral designs, the increase of relative density of cellular design resulted in a decrease of strength, whereas the opposite trend was observed for the BCC designs. As PLA exhibits higher mechanical strength, it is intuitive that an optimal material composition for the highest interfacial strength would consists of larger percentage of PETg and smaller percentage of PLA. However, for the BCC designs, such rather simplistic design approach does not appear to hold valid. Another observation was that the BCC-based interfaces exhibit the highest overall strength among the three designs (Fig.17). For stand-alone cellular structures, BCC typically exhibits relatively low strength, especially when compared to a stretch-dominated design (e.g. octahedral).

For stand-alone cellular structures, BCC typically exhibits relatively low strength, especially when compared to a stretch-dominated design (e.g. octahedral). Therefore, it could be reasonably concluded that the classic design rules of cellular structures are significant to the design of cellular interfaces, however the actual influences could be very different.Fig.10 Fractured Aux-60RD-PETgFig.11 Stress-strain curves for re-entrant auxetic cellular interface samples2252Fig.12 Stress-strain curves for octahedral cellular interface samplesFig.13 Stress-strain curves for BCC cellular interface samplesa. View into the PLA sideb. View into the PETg side2253Fig.14 Fracture surface of Oct-40RD-PLAa. View into the PLA sideb. View into the PETg sideFig.15 Fracture surface of Aux-40RD-PLAa. View into the PLA sideb. View into the PETg sideFig.16 Fracture surface of Aux-50RD-PLAa. Tensile strengthb.

View into the PETg sideFig.15 Fracture surface of Aux-40RD-PLAa. View into the PLA sideb. View into the PETg sideFig.16 Fracture surface of Aux-50RD-PLAa. Tensile strengthb. Elongation at breakFig 17: Tensile properties of cellular interfaces with PLA positive componentsWhen the material assignment flipped, the different interface designs exhibit significant change of behaviors. As shown in Fig.18-20, the stress-strain curves of all three cellular interface designs with PETg positive components exhibit completely different characteristics compared to those with PLA positive components.For the re-entrant auxetic design (Fig.18), the effect of relative density on maximum strength become less significant, while the behaviors of samples for the same design become more consistent. This was found to be mainly attributed to the more consistent failure mechanism with these structures, which is dominated by the weaker PETg phase.

This was found to be mainly attributed to the more consistent failure mechanism with these structures, which is dominated by the weaker PETg phase. With the PETg positive components, the cellular strut features become the critical features of the interface structures, which upon failure exhibit both fracture and pull-out behaviors. In addition, the quality of the PLA2254negative component also appears to be lower compared to the counterparts with PLA as the material of the positive components. This was hypothesized to be caused by the setting with the process parameters but was not further investigated. On the other hand, the stress-strain behaviors of the octahedral designs (Fig.19) become more scattered. It was also found that the fracture surfaces of the octahedral samples exhibit relative smooth tomography.

It was also found that the fracture surfaces of the octahedral samples exhibit relative smooth tomography. The consistency of the relationship between the increased variability of the stress-strain curves with the cellular interfaces and the fracture surface smoothness suggests a potential general design guideline for such interfaces that with increased efficiency of the design of the two component phases, the variability of the resulting structures also become less consistent due to the non-negligible effects of multiple failure mechanisms. Lastly, the effect of relative density design on the overall performance of the BCC interface appears less significant compared to be less significant compared to the other cellular topology designs as well as the counterpart structures with the PLA positive components (Fig.20).

Additional information about the mechanical properties of the three interfaces with PETg positive components is shown in Fig.21.Fig.18 Stress-strain curve of re-entrant auxetic interface with PETg positive componentFig.19 Stress-strain curve of octahedral interface with PETg positive component2255Fig.20 Stress-strain curve of BCC interface with PETg positive componenta. Tensile strengthb. Elongation at breakFig.21: Tensile testing results of PETG+ samplesAn additional FEA analysis was carried out with the three types of interfaces at varying levels of cellular relative density designs. The FEA analysis was set up using perfectly-elastic materials using the elastic modulus obtained from experimentation. Without the consideration of plasticity within the materials, it was not expected that the FEA provides accurate estimation of the behaviors of the structures. However, even for elastic modulus estimation the FEA results differed from the experimentation.

However, even for elastic modulus estimation the FEA results differed from the experimentation. As shown in Fig.22, the re-entrant auxetic interface was predicted to exhibit the highest overall elastic modulus at all relative density levels. Such discrepancy is again likely attributed to the processing-induced issues, as the material extrusion systems used for this study consistently appeared to exhibit quality consistency issues with the cellular interface structures, possibly due to the frequent switching of active material extrusion printheads. This remains a topic for further investigation.2256Fig.22 Elastic modulus of cellular interface samples with PETg positive components4. ConclusionThe current study investigated a new bi-material interface design concept based on cellular structures. Overall, the cellular design-based interface exhibit significantly improved overall ductility and mechanical property controllability.

Overall, the cellular design-based interface exhibit significantly improved overall ductility and mechanical property controllability. The unit cell topology of the cellular structures exhibit significant effects on the behaviors and properties of the cellular interfaces, which could potentially be exploited for property control with these designs. In addition, the design of the interface cellular structures appears to be influence by the mechanical properties of the two component materials. Finally, the behaviors of the cellular interfaces appear to differ from the classic knowledge with cellular structure designs, although some of the design rules with cellular unit cell still appear to be significant with the interface designs.5. AcknowledgementThe authors would acknowledge the support of experimental resources from Additive Manufacturing Institute of Science and Technology (AMIST) Center at University of Louisville.Reference[1][2][3]A. Bandyopadhyay and B.

Bandyopadhyay and B. Heer, "Additive manufacturing of multi-material structures," Materials Science and Engineering: R: Reports, vol. 129, pp. 1-16, 2018. H. Watschke, L. Waalkes, C. Schumacher, and T. Vietor, "Development of novel test specimens for characterization of multi-material parts manufactured by material extrusion," Applied Sciences, vol. 8, no. 8, p. 1220, 2018. C. Yang, X. Tian, T. Liu, Y. Cao, and D. Li, "3D printing for continuous fiber reinforced thermoplastic composites: mechanism and performance," Rapid Prototyping Journal, vol. 23, no. 1, pp. 209-215, 2017.[4] M. A. H. Khondoker, A. Asad, and D. Sameoto, "Printing with mechanically interlocked extrudates using a custom bi-extruder for fused deposition modelling," Rapid Prototyping Journal, vol. 24, no. 6, pp. 921-934, 2018. R. Freund, H. Watschke, J. Heubach, and T.

186, p. 108254, 2020. T. Perets, N. B. Ghedalia-Peled, R. Vago, J. Goldman, A. Shirizly, and E. Aghion, "In vitro behavior of bioactive hybrid implant composed of additively manufactured titanium alloy lattice infiltrated with Mg-based alloy," Materials Science and Engineering: C, vol. 129, p. 112418, 2021.[10] S. Verma, C.-K. Yang, C.-H. Lin, and J. Y. Jeng, "Additive manufacturing of lattice structures for high strength mechanical interlocking of metal and resin during injection molding," Additive Manufacturing, p. 102463, 2021.[11] L. Yang, A. Coriano. A comparative study with the topology design of cellular mechanical interface in a bi-material structure. Presentation. International Solid Freeform Fabrication Symposium, Austin, TX, 2021.2258.

 Experimental Study of Ceramic Dental Porcelain Materials Using A 3D Print (3DP) ProcessShanshan Zhanga, Hadi Miyanajia, Li Yanga, Amir Ali Zandinejadb, J.J.S. Dilipa and Brent Stuckera aDepartment of Industrial Engineering, J.B. Speed School of Engineering bDepartment of Oral Health and Rehabilitation, School of Dentistry University of Louisville, KY, 40292AbstractDental porcelain materials have been used widely in dental restorations such as crowns, veneers and onlays. In this study, a commercial dental-grade porcelain powder (IPS InLine Dentin) was investigated for compatibility with 3D printing. An extensive experimental study was carried out to evaluate the effects of various in-process and post-process parameters on the geometrical accuracies and porosities of the dental porcelain structures, and optimal process parameters were determined that result in homogeneous shrinkage and minimized part distortions.

This study established a practical guideline for the direct fabrication of dental porcelain structures, which enabled further development of this material that focused on performance improvement.1. InstructionTraditionally, ceramic dental restorations produced using processes, such as hot/cold isostatic pressing (HIP/CIP), sintering or slip casting processes (Li and Liao, 1996, Itoh et al. 1994) generally have limited accuracy and capabilities for product customization, which has gradually become the trend. With the generation and development of CAD/CAM methods, these processes have been commonly used for fabricating dental structures due to high accuracy (Denry and Kelly, 2008, Martorelli, 2013, Mohanty, 2013).In the CAD-CAM process, the starting feedstock is usually a block of material that is usually significantly larger than the part to be made, therefore this also creates a considerable amount of waste.

Additive manufacturing (AM) provides a promising method to produce dental restorations (Gibson et al., 2009). In addition to their abilities to fabricate parts directly from the digital model, AM technologies also offer some other potential advantages such as higher geometrical accuracy control, mass customization and rapid production. Currently there exist limited research and development work focused on functional ceramic fabrication with AM, and the use of AM technologies in dental applications is largely focused on the production of prototypes and molds, and the direct manufacturing of metal dental implants (Wu et al., 2001, Azari et al., 2009). However, studies in this area have demonstrated the feasibility of direct manufacturing of high quality ceramic structures using AM. Wang et al. fabricated permanent dental restorations via a micro-extrusion process using slurry materials (Wang et al., 2006).

Wang et al. fabricated permanent dental restorations via a micro-extrusion process using slurry materials (Wang et al., 2006). It was reported that a direct write based printing process can be applied for the fabrication of zirconia dental prostheses (Ebert et al., 2009), and a powder-binder system was used for patient specific implants and scaffolds (Khalyfa et al., 2007, Lam et al., 2002). Compared to other processes, the 3D printing (3DP) binder jetting process possesses various advantages for ceramicprinting. In addition to its good resolution limit and geometrical accuracy, 3DP also has considerably more flexibility for material selection compared to both extrusion and direct sintering processes. Unlike the direct sintering process, 3DP does not involve extensive thermal input, therefore is also less affected by the thermal shock sensitivity of the ceramic materials.

Unlike the direct sintering process, 3DP does not involve extensive thermal input, therefore is also less affected by the thermal shock sensitivity of the ceramic materials. Therefore, the binder jetting based 3DP processes have the potential for direct manufacturing of functional ceramic structures.In this paper, an experimental study was performed to explore the feasibility of using a 3DP process for high accuracy manufacturing of dental ceramic porcelain structures. The goal of this study is to identify the optimal processing and post-processing parameters for high accuracy part fabrication with dental porcelain materials. The relationships between linear shrinkage, porosity and microstructure of the printed dental porcelain structure and various process parameters were investigated through experiments in order to find the proper parameter sets and understand the influence of each parameter on part quality.

Mechanical properties were not investigated in this paper and will be addressed as an ongoing effort.2. Materials Preparation and Fabrication Methods2.1 Powder preparationThe material used in this study is IPS InLine Dentin powder (Manufacturer: Ivoclar Vivadent Corporate, Principality of Liechtenstein), which is a commerical product for artificial dental restorations such as crowns, veneers and onlays.

The chemical composition and relative basic properties of IPS InLine Dentin powder provided by the manufacturer is shown in Table 1 (InLine scientific documentation, 2010).Table 1 Technical data of IPS InLine Dentin powderStandard composition (in wt%)Flexural Strength Chemical solubility Glass transition temperature Coefficient of thermal expansion (25-500°C)SiO2: 59.5 - 65.5, Al2O3: 13.0 - 18.0, K2O: 10.0 - 14.0, Na2O: 4.0 - 8.0, Other oxides: 0.0 - 4.0, Pigments 0.0 - 2.0 80 ± 20 MPa < 100 µg/cm2 585 ± 10 °C 2 firings: 12.60 ± 0.5 10-6 K-1 4 firings: 13.20 ± 0.5 10-6 K-1The original powder exhibits significant aggregation, which made it unable to be spread uniformly in the printing process. From the SEM microscopy, the IPS InLine Dentin particles had irregular shape and different sizes ranging from 0.3-10 μm, as shown in Fig. 1. Laser-based particle size analysis (PSA, Microtrac S3000) shows that the characteristic size of the aggregation was ~36 μm, shown in Fig. 2.

1. Laser-based particle size analysis (PSA, Microtrac S3000) shows that the characteristic size of the aggregation was ~36 μm, shown in Fig. 2. It was found previously that the irregular morphology and large particle size range significantly reduced its flowability (Yang et al., 2013). In order to improve the flowability of the powder, a flow agent was added. The function of the flow agent is analogous to the addition of sand between two surfaces, which serves as low friction contact media and therefore reduces the resistance of relative motions between the powder particles. The chemical composition of theflow agent is largely irrelevant to its primary functionality, therefore flow agent with composition similar to the original powder was selected.25µm5µmFig. 1 Morphology of original IPS InLine Dentin powder in SEMThe flow agent chosen was Aerosil R 972 Hydrophobic fumed silica powder (COSMOS Plastic & Chemicals), with an average particle size of 16 nm.

1 Morphology of original IPS InLine Dentin powder in SEMThe flow agent chosen was Aerosil R 972 Hydrophobic fumed silica powder (COSMOS Plastic & Chemicals), with an average particle size of 16 nm. This powder is composed of 99.8% fumed silica after treated with dimethyldichlorosilane (DDS) based on a hydrophilic fumed silica with a specific surface area of 130 m2/g (Aerosil R 972 MSDS). Up to 10% volume percentage flow agent was measured and added to the original IPS InLine Dentin powder, and the container with the mixture was shaken by hand until well-proportioned mixed powder was visually discernable. The flowability was evaluated using both angle of repose and particle size analyzer for the original powder and treated powder with flow agent additive.(a) Original Powder(b) Flow agent addedFig.2 Particle size analysis results of the IPS InLine Dentin powderFrom the results of particle size analysis, flow agent has a significant effect on aggregation. In Fig.

In Fig. 2(a), it can be seen that the mean particle size of the original powder is about 25 μm, and the 95 percentile particle size is 52.89 μm. After adding flow agent, the average particle size reduced to about 19 μm as shown in Fig.2(b), and the 95 percentile particle size also decreased to 36.79 μm.The angle of repose test also showed significant improvement of flowability with the prepared powder. The comparison of angles of repose of the original powder and the one with flow agent additive is shown in Fig. 3, and the measurement results are listed in Table 2. Apparently, the addition of 10% by volume flow agent significantly reduced the aggregation and flowability issue of the original powder. This was further verified with a preliminary trial, and was consequently used as the standard powder preparation method for the study.(a) Original (b) Flow agent addedFig.

This was further verified with a preliminary trial, and was consequently used as the standard powder preparation method for the study.(a) Original (b) Flow agent addedFig. 3 Angles of repose of original powder and flow agent-added powderTable 2 Angles of repose of powderPowder type Original 10% (vt.) addition of flow agentAngle of repose (degree) 67 58The ExOne M-Lab system was used for this study, which is a powder bed based binder jetting 3DP process. During the process, binder liquid infiltrates into the powder bed and bonds the powder particles to create the geometries. After the printing operation of each layer, the powderbed is heated by an infrared heater for a set amount of time in order to partially cure the binder and to gain necessary strength. For each new layer, the powder is fed from the powder feeding box via a roller. The procedures are repeated until the parts are completed. The schematic of the ExOne 3DP process is shown in Fig.

For each new layer, the powder is fed from the powder feeding box via a roller. The procedures are repeated until the parts are completed. The schematic of the ExOne 3DP process is shown in Fig. 4.Binder Print HeadParts Building BoxPowder Feeding BoxZXYFig. 4 Schematic of 3D printing process for an ExOne M-lab systemThe M-Lab system uses two special parameters for the control of binder amount and drying level, which are the power level and the saturation level. The default setting for these two parameters are 60% power level and 70% saturation, which was developed for stainless steel powder. In order to identify suitable parameters for the dental porcelain powder used in this study, a preliminary experiment was performed. Two sets of cubic specimens with size of 10x10x10 mm were designed and fabricated under different power levels and saturation levels.

Two sets of cubic specimens with size of 10x10x10 mm were designed and fabricated under different power levels and saturation levels. The fabricated green parts were put into a drying oven and baked at 200°C for 2 hours to fully cure the binder.2.2 Saturation levelThe first set of experiments evaluated the effect of saturation level on the print quality of the porcelain powder. The binder saturation levels were set at 45%, 50%, 60% and 70%, respectively, with the power level fixed at 60%. The printed samples are shown in Fig. 5. It is obvious that saturation level could significantly affect the geometry of green parts. When the saturation level was set as 70%, the printed parts deformed in the direction of roller spreading. This was likely caused by insufficient drying of binder and the resulting low strength of the printed area, which could be displaced under the frictional force of the roller when a new layer of powder is deployed.

As the saturation level reduced, less distortion was observed. On the other hand, when the saturation continued to reduce to 45%, the amount of binder became insufficient, and the strength of green parts was so low that the parts could not be handled without damage. A quick loading test was performed on the green parts to evaluate the maximum compressive force that could be applied to these parts without significant damage. A flat plate was placed on the top of the samples, and weight was gradually added to the plate until the sample crumbled. The results are shown in Fig. 6, and it was apparent that when fully cured, the strength of the green part increases with the increase of binder amount. When the saturation level reduces from 50% to 45%, there was a ~25% reduction of green part strength, which supported the experimental observation.(a)(b)(c)(d)Fig.

When the saturation level reduces from 50% to 45%, there was a ~25% reduction of green part strength, which supported the experimental observation.(a)(b)(c)(d)Fig. 5 Green parts printed using different saturation levels, (a) 45%; (b) 50%; (c) 60%; (d) 70%In order to analyze the effect of saturation level on the quality of final parts after sintering, the same set of green parts were sintered at 850°C for 30 minutes. Specimens of each saturation level were sectioned at mid-plane along the vertical direction and polished for optical microscopy (Olympus MX51). Pore distributions were observed under microscopy, as shown in Fig. 7 with 200x magnification. From the microscopy, the size of pores of samples with 45% saturation level ranged from 5- 150 µm, while the pores in samples with 70% saturation level were considerably smaller with a maximum size of ~45 µm.

The maximum pore diameters become smaller as the saturation level increases, but the total porosity does not seem to change significantly. One possible cause of the porosity increase at lower saturation level is the surface void formation during the powder deposition by the roller, which results in the lower binding strength. The addition of the flow agent might also contribute to the porosity, as the flow agent particles might result in lower binder wettability due to the hydrophobic surface. Based on the analysis of porosity in different saturation levels, 50% saturation was chosen for further study since it yields adequate handling strength and minimizes in-process part distortions.Load Tests on Effect of Saturation100.00090.000) a P k ( e r u s s e r P80.00070.00060.00051.71565.20084.59388.22050.00040.00045%50% Saturation60%70%Fig.

6 Maximum compressive load of green part under different saturation levels(a)(b)200µm200µm(c)(d)200µm200µmFig. 7 Microstructure of final parts with different saturations (heating treatment at 850°C for 30 min), (a) 45%; (b) 50%; (c) 60%; (d) 70%2.3 Power LevelSince the powder bed is cured after each layer is printed, the power level could potentially have significant influence on the in-process part qualities. Therefore, in the second set of experiments, the power levels were set as 55%, 60% and 65% while the saturation levels were kept at 50%. The printed samples are shown in Fig. 8. From the measurement results, the power level does not seem to have apparent effects on the geometrical accuracies of the green parts. The same set of samples was also sintered at 850°C for 30 minutes and prepared in the same way for microscopy. The loading test of the green parts at all three power levels yield adequate strength as shown in Fig.

The loading test of the green parts at all three power levels yield adequate strength as shown in Fig. 9, while higher power levels appeared to be more advantageous. On the other hand, higher power level corresponded to an increase of drying time, which was significant due to the fact that the powder bed was dried at every layer. The microstructure of the samples is shown in Fig. 10, and there was also no significant dependence between the power level and the porosity of the final parts. This might be explained by the relatively low saturation level used for these studies. It could be reasonably assumed that the binder could be adequately cured at the range of power levels experimented with(e.g. 55%-65%), therefore no significant difference could be observed on the green parts.

55%-65%), therefore no significant difference could be observed on the green parts. It could also be expected that at higher saturation levels, higher power level might be needed to achieve satisfactory handling strength.Therefore, based on the preliminary experimental results, 60% power level was selected asthe standard parameter for further study.(a)(b)(c)Fig. 8 Green parts printed in different power Levels, (a) 55%; (b) 60%; (c) 65%) a P k ( e r u s s e r PLoad Tests on Effect of Power Level 94.383100.000 90.000 80.000 70.000 60.000 50.000 40.00065.20062.55255%60%65%Power LevelFig. 9 Effect of power level on the compressive strength of the green parts(a)(b)(c)200µm200µmFig. 10 Microstructure of final parts with different power levels (heating treatment at 850°C for 30 min), (a) 55%; (b) 60%; (c) 65%3.

10 Microstructure of final parts with different power levels (heating treatment at 850°C for 30 min), (a) 55%; (b) 60%; (c) 65%3. Sintering ExperimentsAfter the printed green parts are fully cured in the drying oven at 200°C for 2 hours, they need to be cleaned and sintered to achieve the desired strength and density. A general firing schedule for dental porcelain ceramic is shown in Fig. 11. The firing process started at room temperature, and heated at a set rate to 500°C for half an hour (phase I), which aimed for burning out the binder completely. After the burnout step, the part continued to be heated to the target sintering temperature and subsequently held for a set amount of time (phase II), and then cooled at 10°C/min to room temperature. Similar to the sintering of metal powders and other ceramic powders, the sintering temperature, holding time and heating rate could have significant effect of200µmthe shrinkage, microstructure and distortion of the final parts.

Therefore these parameters were investigated in this study in details. The experimental design is shown in Table 2, and the levels of each variable were determined by standard firing schedules used in dental clinics for dental porcelain.T/°CⅡⅠFiring schedule:t/hⅠ: 500°C × 30min; Ⅱ: T × t; Cooling rate: 10°C/minFig. 11 Firing scheduleTable 2 Factors and variables of experimentFactors Temperature Holding time Heating rateLevels 750, 800, 850, 900, 925, 950°C Instantaneous (1min), 10min, 30min, 2h, 10h 100°C/h, 500°C/h, 5000°C/hSintering of ceramics is a complicated process that involves mass transport driven by multiple mechanisms. For low temperature sintering, the mass transporta is mainly controlled by the surface energy and/or interface energy of the ceramic powder. From an atomic diffusion perspective, higher firing temperature or longer holding time would facilitate greater diffusion and therefore promote the sintering process.

From an atomic diffusion perspective, higher firing temperature or longer holding time would facilitate greater diffusion and therefore promote the sintering process. In the initial sintering stage, the driving force fordiffusion is larger due to the large surface areas. However, as the sintering develops, excessive temperature or holding time might cause excessive grain growth, and the overall properties of sintered parts would deteriorate.During the first stages of sintering, diffusion occurs mainly through the sintering neck between two contacting particles, which results in shrinkage of the pores and the growing of the neck size between the particles. This in turn causes shrinkage of the part. As diffusion continues, the adjacent grain boundary intersects and forms networks. Driven by surface tension, the grain boundary migrates, and the grains begin to grow.

As diffusion continues, the adjacent grain boundary intersects and forms networks. Driven by surface tension, the grain boundary migrates, and the grains begin to grow. From powder metallurgy theory, the sintering could be roughly estimated by the growth of the sintering neck (R) as a function of sintering temperature (T) and holding time (t).−11 7⁄𝑇⁄ 𝑡 (1) From Eq.(1) (Huang, 1997), it is apparent that temperature has a pronounced effect on sintering. Therefore, a multi-step experimental design approach was adopted, which was expected to be more efficient and accurate compared to a Taguchi design or full scale design. The multi- step experimental design is demonstrated in detail below.𝑅 ∝ 𝑒Step 1: Firing temperatureThe effect of fire temperature on the density and microstructure of the samples was investigated first, since temperature is the most influential factor on sintering. Table 3 shows the details of this experiment.

From Eq.(1), the sintering parameters of 850°CX30min was selected as the baseline parameters, a sintering index could be obtained. Experiments were then carried out by considering ±50% and ±20% variation of the sintering index, which was achieved by changing the holding time.

After an optimal group of temperature and time was obtained from Step 1 and 2, the effect of heating rate was investigated at three levels from slow heating to very rapid heating, as shown in Table 5.#17Initial temperature heating rateRT100°C/hTable 5 Heating rate schedule heating ratePhase Ⅰ 500°CX30min100°C/hphase Ⅱ 850°CX30mincooling rate600°C/h10RT500°C/h500°CX30min500°C/h850°CX30min600°C/h18RT5000°C/h500°CX30min5000°C/h850°CX30min600°C/hThroughout the experiments, the linear shrinkage values were evaluated for each sample,which was calculated according to ASTM C326-09 as:𝑆 = [(𝐿1 − 𝐿2)/ 𝐿1] × 100%where S is percentage linear dimensional change, 𝐿1 is average length, width, or thickness of specimen before sintering, and 𝐿2 is average length, width, or thickness of specimen after sintering.

After the parts were sintered, the porosity of each group of specimens was measured using the Archimedes method (Spierings, 2011) according to ASTM B962-08 as:𝑅𝐷 = 1 − 𝜌𝑒 𝜌𝑛⁄where RD is porosity, 𝜌𝑒 is measured density, and 𝜌𝑛 is nominal density.4. Results and Discussion4.1 4.1.1 Effect of heating temperatureLinear shrinkageFig. 12 shows the linear shrinkage of the samples sintered at different temperatures for 1 minute. As sintering temperature increased, the linear shrinkage in horizontal directions (X and Y directions) on the top surface of the sintered parts showed an ascending trend until 850°C, and then remained largely constant at about 25%. However, the curve of the linear shrinkage in horizontal directions at the bottom surface showed an obvious decrease after a small peak of linear shrinkage at 850°C. It could be speculated that the glassy phase has a lower melting point, which led to reduction in porosity and enhanced densification of the body.

It could be speculated that the glassy phase has a lower melting point, which led to reduction in porosity and enhanced densification of the body. On the other hand, at elevated temperature the glass phase might partially collapse under its own gravity due to the loss of strength. This also explains the shrinkage phenomenon in the vertical direction (Z direction), which showed an on-going ascending trend as the temperature increases, with the maximum value up to 32% (at 950°C).In the group of specimens which were sintered for 30 minutes (Fig. 13), the linear shrinkage in horizontal directions on the top surface kept increasing after 850°C. When the temperature increased to the next level (925°C), the sintered body showed apparent signs of melting. At 950°C, the specimen completely lost its original cubic shape due to melting.

When the temperature increased to the next level (925°C), the sintered body showed apparent signs of melting. At 950°C, the specimen completely lost its original cubic shape due to melting. The linear shrinkage on the bottom surface appeared to reach maximum at 850°C before it started to decrease, which was similar to the group sintered for 1 minute.(2)(3)Effect of Temperature (Holding for 1 min)35.00%30.00%25.00%20.00%15.00%750°C800°C850°C900°C925°C950°CTop XTop YBottom XBottom YZFig. 12 Trend of linear shrinkage in X, Y and Z building directions as temperature changed (holding for 1 min)Effect of Temperature (Holding for 30 min)33.00% 31.00% 29.00% 27.00% 25.00% 23.00% 21.00% 19.00% 17.00% 15.00%750°C800°C850°C900°CTop XTop YBottom XBottom YZFig.

13 Trend of linear shrinkage in X, Y and Z building directions as temperature changed (holding for 30 min)4.1.2 Effect of Holding TimeFor the group sintered at 850°C, the linear shrinkage showed a slight increasing trend with longer sintering time until the holding time reached 2 hours, and the linear shrinkage started to decrease on the top surface of horizontal directions with longer holding time (Fig. 14). The linear shrinkage at the bottom of horizontal directions also had the same trend except that the maximum appeared at the holding time of 30 min. On the other hand, the linear shrinkage in the vertical direction exhibited more or less constant linear shrinkage. The apparent decrease of linear shrinkage at prolonged sintering times might be explained by the sintering dynamics. When sintering develops to the late stage in which the pores are completely enclosed and become spherical, further sintering no longer effectively eliminates the porosity.

When sintering develops to the late stage in which the pores are completely enclosed and become spherical, further sintering no longer effectively eliminates the porosity. Instead, under the thermal diffusion effect, the pores could start to migrate and coalesce, which could potentially lead to macro-scale distortion of the sintered structures.Effect of Holding Time in Phase II (850°C)29.00%27.00%25.00%23.00%21.00%19.00%17.00%1min10min30min2h8.6h10hTop XTop YBottom XBottom YZFig. 14 Trend of linear shrinkage in X, Y and Z building directions as holding time changed (sintering at 850°C)For the group of specimens sintered at 900°C as shown in Fig. 15, excessive sintering was apparent beyond 1 minute of holding time. This was also clearly shown in the shapes of the final parts, as partial melting and part distortion became obvious at longer holding times (Fig.

This was also clearly shown in the shapes of the final parts, as partial melting and part distortion became obvious at longer holding times (Fig. 16).Effect of Holding Time in Phase II (900°C)32.00%29.00%26.00%23.00%20.00%17.00%1min10min30minTop XTop YBottom XBottom YZFig. 15 Trend of linear shrinkage in X, Y and Z building directions as holding time changed (sintering at 900°C)Instantaneous (1min)10min30min2hFig. 16 Specimens sintered at 900°C holding for different time4.1.3 Effect of Heating RateThe heating rate experiment was performed at 900°C for 1 minute, and the results of the linear shrinkage are shown in Fig. 17. The linear shrinkage in all directions showed the same trend as the heating rate changed from 100 °C/h to 5000 °C/h.

17. The linear shrinkage in all directions showed the same trend as the heating rate changed from 100 °C/h to 5000 °C/h. Generally, the vertical direction had larger shrinkage than horizontal directions, and top surface also had a little higher values than the bottom surface.For very rapid heating rate (5000 °C/h), smaller linear shrinkage could be expected due to the insufficient sintering. On the other hand, at very slow heating rate (100 °C/h), the samples were in fact subject to prolonged holding time at high temperature ranges, which might effectively cause gravity induced distortion and over-sintering distortion as discussed before.Effect of Heating Rate (900°Cx1min)27.00%26.00%25.00%24.00%23.00%22.00%21.00%20.00%100°C/h500°C/h5000°C/hTop XTop YBottom XBottom YZFig. 17 Linear shrinkage in X, Y and Z building directions at different heating rates4.2 4.2.1 Effect of heating temperaturePorosityAs shown in Fig.

17 Linear shrinkage in X, Y and Z building directions at different heating rates4.2 4.2.1 Effect of heating temperaturePorosityAs shown in Fig. 18, as the sintering temperature increases, the porosity of the final part reduces steadily due to more sufficient sintering. Also, at 925°C/30min significant melting already occurred, therefore it appeared that the minimum porosity attainable porosity with the current process method is about 7% for the dental porcelain.0.1200.110Effect of heating temperature (1min)0.109Effect of heating temperature (30min)0.1000.094y t i s o r o P0.1000.0900.0800.0700.0890.0820.0650.0680.072y t i s o r o P0.0900.0800.0700.0600.0800.0770.0660.0600.0500.050750°C800°C900°C Temperature850°C925°C950°C0.040750°C800°C850°CTemperature900°CFig. 18 Porosity effect of heating temperature4.2.2 Effect of holding timeAs shown in Fig.

18 Porosity effect of heating temperature4.2.2 Effect of holding timeAs shown in Fig. 19, the holding time does not appear to have significant effect on the porosity at both 850 °C and 900 °C. When the holding time was shorter than 2 hours, the porosity did not change significantly at different holding times. However when the holding time was longer than 2 hours, the porosity values showed apparent decrease. This is in agreement with the linear shrinkage observed at these temperature levels.

However when the holding time was longer than 2 hours, the porosity values showed apparent decrease. This is in agreement with the linear shrinkage observed at these temperature levels. This observation also implies that when sintered at proper sintering temperatures, the holding time of the parts could be shortened considerably without significantly affecting the overall porosity, which is very useful for reducing the total production time for dental ceramic structures.Effect of holding time (850°C)Effect of holding time (900°C)0.1000.100y t i s o r o P0.0900.0800.0700.0600.0820.0790.0770.0810.0720.055y t i s o r o P0.0900.0800.0700.0600.0650.0680.0660.0660.0500.0501min 10min 30min2h8.6h10h1min10min30min2hTimeTimeFig. 19 Porosity effect of holding time4.2.3 Effect of heating rateThere was no significant effect of heating rate on porosity as shown in Fig.

19 Porosity effect of holding time4.2.3 Effect of heating rateThere was no significant effect of heating rate on porosity as shown in Fig. 20, which was unexpected since the linear shrinkage of samples at 5000 °C/h heating rate was significantly lower. A possible reason could be that those samples exhibit some through-porosity due to the insufficient sintering, which resulted in water infiltration into the interior of the samples during the weight measurement. Further inspection might be needed to clarify this phenomenon.Effect of heating rate (900°Cx1min)0.1000.090y t i s o r o P0.0800.0700.0600.0620.0650.0650.0500.040100°C/h500°C/h5000°C/hHeating RateFig. 20 Porosity effect of heating rate4.3 MicrostructureFrom the results of 4.1 and 4.2, the linear shrinkage and porosity could keep a relative stability at 850°C for 1-30min and 900°C for 1-10min, which is also illustrated by the microscopyof these samples as shown in Fig. 21.

21. In general, as sintering develops, the pore morphology and size distribution should become more homogenous, which could also be seen from the microscopy. The microscopy from the heating rate experiment samples are also indicative, as larger porosity and pore size distribution as a result of insufficient sintering could be clearly seen from Fig. 22.850°C x 1min900°C x 1min200µm200µm850°C x 10min900°C x 10min200µm200µm850°C x 30min200µmFig. 21 Images of distribution of pores in optical microscopy (Shrinkage Stable)100°C/h500°C/h5000°C/h200µm200µmFig. 22 Images of distribution of pores (Heating rate)200µm5. Dental crown printingA dental crown is a tooth-shaped cap that is placed over a tooth to cover the tooth and improve its appearance. Thin wall features are designed in side of crown structures.

Dental crown printingA dental crown is a tooth-shaped cap that is placed over a tooth to cover the tooth and improve its appearance. Thin wall features are designed in side of crown structures. Based on the experimental results obtained from the study, a batch of dental ceramic crown prosthese from a scanned model were printed using the ExOne M-Lab system using the optimal method with relatively stable linear shrinkage. The 3D model was rescaled 40% larger than the original scanned model to obtainthe correct size after shrinkage. The rescaled 3D model of a dental crown prosthesis is shown in Fig. 23. The linear shrinkages in each direction (X, Y and Z directions) in the dental crown were studied. Fig. 24 shows the final parts that were printed using processing parameters of 50% saturation and 60% power level, and sintered at 900°C for 1 minute.Fig. 23 3D digital model of a dental crown prosthesisFig.

23 3D digital model of a dental crown prosthesisFig. 24 Porcelain crowns by 3D printing processThe average linear shrinkage values in each of three directions of the fabricated dental crowns are shown in Fig. 25. The linear shrinkage in the Z (vertical) direction was larger than X and Y (horizontal) directions. The lateral shrinkage values were similar to the experimental results obtained from the cubic specimens. However, the linear shrinkage value in the Z direction was 15% larger than the results from the cubic specimens. This might be caused by the partial collapse of the structure during sintering. Since the crown prosthesis has thin wall features, when the green parts were placed in the furnace upside down, slight gravity induced distortion might have occurred at the crown surfaces.

Since the crown prosthesis has thin wall features, when the green parts were placed in the furnace upside down, slight gravity induced distortion might have occurred at the crown surfaces. Furthermore, with Archimedes method, the porosity was measured to be an average value of 5.7%, which is in good agreement with the cubic specimens.Dimensional linear shrinkage35.00%30.69%30.00%25.00%24.78%24.47%20.00%15.00%10.00%5.00%0.00%XYZFig. 25 Linear shrinkage in X, Y and Z directionsFor comparison purposes, a cubic specimen with the same dimensions was also fabricated with the standard method currently used in dentistry. The specimen was made by manually forming the cubic shape from the wetted InLine porcelain powder, which was then sintered using the standard firing schedule (900 °C for 1 minute, heating rate 500 °C/hour). Fig. 26 shows the sintered specimen.

Fig. 26 shows the sintered specimen. The geometrical accuracy of the fabricated specimen was not very good, which might be caused by the manual manipulation during the fabrication. The porosity of this specimen was 3.3%, which was lower than the ones fabricated using the 3DP process. Fig. 27 shows the distribution of pores under optical microscopy, and the image confirmed its lower porosity.Fig. 26 A cubic specimen fabricated by a standard manual method200µmFig. 27 Microstructure of the cube specimen fabricated by a standard manual method6. Summary and ConclusionAs one of the less well understood types of AM processes, the 3DP process has a number of control parameters that can significantly affect the fabrication quality. The experimental study performed in this paper successfully revealed the relationships between these parameters and the geometrical accuracy and microstructure of the final parts for a dental porcelain ceramic powder.

More specifically, some of the conclusions from this study can be directly applied to further development with this material:1) Flow agent had a significant effect on aggregation reduction, and improved flowability significantly for the printing process.2) When saturation was set above 60%, the green parts began to deform under the friction of the roller. But when saturation was as low as 45%, the binder could not provide enough strength to bind the powder. Additionally, saturation had a significant effect on the pores of sintered parts in micrographs. The pore diameters decreased as the saturation level increased, while total porosity did not seem to be significantly affected.3) Power level does not have a significant effect on geometry and microstructure of the green and final parts.4) Effect of temperature: When distortion does not occur, linear shrinkage in horizontaldirections increases as the temperature increases.

At higher temperature, parts are more likely to exhibit gravity induced distortion, therefore reducing the accuracy of the part. 5) Effect of holding time: Holding time has a very mild effect on the linear shrinkage and porosity of the parts. At higher temperature levels, longer holding time could result in significant distortion of the parts. On the other hand, the surface finish could be improved with longer holding time.6) Effect of heating rate: Overall heating rate has a significant effect on the linear shrinkage of the parts, which is primarily caused by the amount of sintering at different heating rates.

At very rapid heating rate (5000°C/h), longer holding time would be necessary in order to avoid insufficient sintering.7) The optimum process parameters obtained from the cubic specimens were successfully applied to a thin-walled dental crown prosthesis, which exhibited good agreement of linear shrinkage values in all three directions.8) Binder jetting 3DP processes can produce better dimensional accuracies compared to the current standard fabrication methods. However, the higher porosity level in the final parts might affect the mechanical performance of the dental ceramic structures produced by 3DP, which needs to be characterized in further studies.AcknowledgementThe authors gratefully acknowledge the Rapid Prototyping Center (RPC) for their support of facilities, test equipment and other assistance throughout this project, and Dr. Samuel Dilip for the ceramic sample microscopy.ReferenceJ. Li, H. Liao and L.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperAn Experimental Study of Design Strategies for Stiffening Thin Plates under CompressionIrving E. Ramirez Chavez, Cameron Noe, Vigneshwaran Sekar, Shainil Jogani, Siddharth Israni, Dhruv Bhate* Ira A. Fulton Schools of Engineering, Arizona State University, Tempe, AZ 85287 *Corresponding Author: dhruv.bhate@asu.edu (D. Bhate)AbstractIncreasing stiffness and failure loads while minimizing mass is useful in many engineering applications, including the design of thin plates and shells. In this paper, the performance of thin plates using a range of stiffening approaches were studied for the specific instance of compressive loading. Periodic, graded, stepped, “Voronoi” stochastic, and topologically optimized patterns were explored.

Periodic, graded, stepped, “Voronoi” stochastic, and topologically optimized patterns were explored. These stiffening designs were realized using different software tools and manufactured with the Selective Laser Sintering (SLS) process. These 3D printed specimens were tested under compression to assess their mechanical response. Videos of these tests were recorded to study the shape of the failure modes. This data was analyzed to determine the performance of the different stiffener designs, in comparison to the performance of baseline plates without any stiffening.

This data was analyzed to determine the performance of the different stiffener designs, in comparison to the performance of baseline plates without any stiffening. The study concludes with a discussion of the results and their implications for stiffening thin plates, showing that triangular and stochastic stiffening strategies show particular promise in increasing specific compressive stiffness and specific buckling load.Keywords: Design for Additive Manufacturing, Stiffener, Compression, Plate, Shell, Topology OptimizationIntroductionLoad bearing components in large, planar equipment such as housings, pressure vessels and ribs, often have horizontal and/or vertical stiffeners distributed throughout their surface. These stiffeners play a major role in determining the mechanical performance of the structures.

These stiffeners play a major role in determining the mechanical performance of the structures. There are several ways plates can be stiffened – like all cellular materials, the key questions involve selecting a unit cell, its size and distribution, the associated design parameters, and finally integration of these units into the larger structure [1]. Conventionally, stiffener patterns are derived from repetition of primitive shapes such as circles, squares and triangles. Stochastic and biomimetic patterns have also been explored. Chacón et al. [2] stiffened a steel plate with longitudinal and transverse stiffeners and concluded that both increased the resistance of the steel plate to loading when compared to unstiffened members. Xu et al. [3] studied the elastic stability of horizontally stiffened steel plate walls in compression and found that closed section stiffeners can effectively increase the elastic critical buckling stress. Vu et al.

Vu et al. [4] investigated the optimum locations of multiple longitudinal web stiffeners for steel plates under bending by using a gradient-based interior point optimization algorithm and suggested an efficient method for optimizing the location and the flexural rigidity of longitudinal stiffeners installed at one side of the web of plates. Eswara Kumar et al. [5] explored various stiffener patterns (honeycomb, triangular, square, semi-circular, rectangular, circular, helical etc.), and concluded that the helical stiffener design was optimum for the designs evaluated. Glassman et al. [6] examined the effectiveness of stiffeners for enhancing the shear capacity of slender plates at ambient and elevated temperatures and found that for lower1893temperatures, a stiffener oriented along the compression diagonal provides the most improvement in shear strength, but for temperatures above 600°C the stiffener is ineffective regardless of the stiffener’s geometric orientation. Ali et al.

Ali et al. [7] studied the bending of biomimetic scale covered beams under discrete non-periodic engagement and made a comparative study between the symmetric and non-symmetric structures. Putra et al. [8] carried out structural optimization of stiffener layout for a plate by using number and type of stiffeners as discrete design variables and optimized their designs using a genetic algorithm.In this experimental study, different plate stiffening approaches are examined under compressive loading. Of specific interest is how periodic, graded, and stochastic patterns influence the compressive stiffness and buckling load, as a function of mass (“specific” properties), and how these compare to unstiffened and topologically optimized baselines. In choosing these three approaches, we follow the classification of cellular materials proposed by Bhate et al. [9].

In choosing these three approaches, we follow the classification of cellular materials proposed by Bhate et al. [9]. The relative merits of these different approaches to stiffening plates have received limited study in the literature, particularly in the context of design optimization and Additive Manufacturing (AM). In the next section, the methods employed in this study are discussed, ranging from design, manufacturing and testing under compression. Results are presented, followed by an analysis in terms of the metrics of interest (specific compressive stiffness, and specific buckling load).

Results are presented, followed by an analysis in terms of the metrics of interest (specific compressive stiffness, and specific buckling load). The paper concludes with an interpretation of these findings, and their implications for future work.MethodsThe approach used in this study is summarized in Figure 1 and can be broken down into three different steps: (i) the CAD design of the specimens to be tested, conducting using three different software packages, (ii) manufacturing these specimens using the Selective Laser Sintering (SLS) process with Nylon 12 (PolyAmide 2200), and (iii) compression testing conducted on a mechanical test frame. In this section, each of these methods is discussed in more detail.Specimen DesignManufacturing.,,.TestingSolidWorks, Fusion 360 nTopology Element SolidThinking InspireEOS Formiga P110 Nylon 12 (PA 2200)/NSTRON 8801 Compression, 10-3 s·1Figure 1.

The three key steps of this work involved specimen design, manufacturing and testing, along with the associated software and hardware utilized.1894i. Specimen DesignFive families of plate design concepts were explored in this work: unstiffened baseline,square, triangular, stochastic and topology optimized. Each of these is discussed in turn.a. Unstiffened baselineIn order to be able to compare the different stiffening strategies, a baseline was first generated with dimensions shown in Figure 2, using SolidWorks CAD software. An initial design exploration study was conducted to establish a minimum wall thickness that could be reliably printed with the SLS process (at the lower bound for manufacturability) and gave good test results under compression (to set the upper and lower bounds for testability). Specimens were designed between 1 and 6 mm thick to conduct this study.30.00I o 0 0 1Figure 2.

Specimens were designed between 1 and 6 mm thick to conduct this study.30.00I o 0 0 1Figure 2. Solid Plate CAD MeasurementsCompression tests showed good stability for the 1.5, 2, 2.5, 3, 4 and 6 mm thickness plates. The 1mm plate failed to generate usable data, likely due to the fact that the force needed to deform it was very low in comparison to the sensitivity of the compression tester used. As a result of these evaluations, solid plates with thicknesses 1.5 and 2mm were designed as the baselines for the rest of the study.A second exploration study was conducted to establish the extrusion thickness of the stiffeners in both sides of the specimen. Different extrusion thicknesses and widths were printed in order to select the one with the best resolution as determined visually. The selected values were .5 mm for the thickness of the beam and 2 mm for the extruded length, the latter shown in Figure 3.1895I __,.._Figure 3.

The selected values were .5 mm for the thickness of the beam and 2 mm for the extruded length, the latter shown in Figure 3.1895I __,.._Figure 3. Solid plate CAD file showing an extrusion thickness of 2mm on a 1.5mm thick plateb. TriangularTriangular patterns are commonly used for plate stiffening. In this work, we used two different triangular unit cell designs with a periodic and graded approach, in five different designs, shown in Figure 4. Designs 1 and 2 represent periodic triangular patterns with a cross- and diagonal layout, respectively. Designs 3, 4 and 5 are graded, in an attempt to study its effects on buckling (i.e. change specific bucking load). Designs 3 and 4 have the diagonal layout, design 5 has the crossed layout.12345Periodic Triangular Pattern IPeriodic Triangular Pattern 2Graded Triangular Pattern IGraded Triangular Pattern 2Graded Triangular Pattern 3Figure 4.

The five different triangular patterns evaluated in this work: designs 1 and 2 are periodic, designs 3-5 are graded.c. SquareThe design software Fusion360 [10] was used to develop three square-based stiffening patterns, shown in Figure 5, as a counterpoint to the triangular patterns in Figure 4. Once again, graded structures were implemented to ascertain their effects on the specific buckling load.1896678l .I _. ::a Jo 1.11,111111 1111111111,Ji,I. "II :&I !I 11 'I I i1I11t111JIII I I I l I I I I .1 l 1111111111 I .l ,J .l .I .I .1 .1 ,I IGraded Square Pattern l Figure 5. The three different square patterns evaluated in this work: designs 6 is periodic, designs 7 and 8 are graded.Periodic Square PatternGraded Square Pattern 2d. StochasticFour types of stochastic stiffening strategies were implemented based on the concept of “Voronoi diagram” [11], using the software package Element from nTopology [12], as shown in Figure 6.

Element allows the prescription of a uniform stochastic distribution of a certain average cell size, used in Design 12. For Designs 9-10, a step function is employed, and a drop-off graph with a 50mm range of influence is used for design 11, resulting in varying gradations and densities.9101112Step Stochastic IStep Stochastic IIGraded StochasticUniform StochasticFigure 6. Four different stochastic patterns evaluated in this work: designs 9-10 employ a step function, design 11 is graded, design 12 is uniform.e. Topology OptimizedFinally, as a contrast to the cellular based stiffening approaches discussed above, three specimens were designed using topology optimization using the Inspire software from SolidThinking [13]. Starting from a 2mm thick base, plate 2 mm in each direction extending out of each face of the plate was designated as a design space for adding mass during the topology optimization.

Starting from a 2mm thick base, plate 2 mm in each direction extending out of each face of the plate was designated as a design space for adding mass during the topology optimization. The base of the I-beam was fixed and the top loaded with a force of 65.59 psi. This force value was chosen since it was close to the the failure load of comparable solid plate test specimens established during experimental testing, discussed in a later section. For these specimens the objective was to maximize stiffness. The amount of remaining mass can be selected1897before the optimization is executed. For this study, three different samples were created: 10% mass retention, 30% mass retention, and 50% mass retention. No safety factors were specified.13141530% Mass Retention TopOpt Figure 6.

No safety factors were specified.13141530% Mass Retention TopOpt Figure 6. Topology optimization approach to stiffening plates with three different mass retention constraints specified with the objective of maximizing stiffness10% Mass Retention TopOpt50% Mass Retention TopOptii. ManufacturingAll specimens were printed on the EOS Formiga P110 Selective Laser Sintering (SLS) printer, shown in Figure 7a. PA 2200 Nylon powder was used, with material properties (for topology optimization studies discussed previously, and for mass calculations) extracted from online datasheets for printed PA 2200 Nylon powder from EOS [14]. Specimens were printed in the same relative orientation, with their plate faces parallel to the printer’s Z-axis, as shown in Figure 7b. Specimens were printed in two different, but successive print batches, but due to the small number of specimens in this study, with no replications, batch effect was not studied.

Specimens were printed in two different, but successive print batches, but due to the small number of specimens in this study, with no replications, batch effect was not studied. Parts were blasted with compressed air and rinsed with water and dried, to remove loose nylon powder.FOAMIGA P 110(a) (b) Figure 7 (a) EOS Formiga P110 SLS system used in this study, (b) 3D nested layoutused to fabricate these specimens, showing orientation of I-beams1898jiii. Compression TestingThe stiffened plates were tested under compression testing at a strain rate of 10-3 s-1 in an Instron 8801 machine with a 50kN capacity, following strain rates recommended in ISO standards for cellular material testing [15]. The main objectives were to find the load value corresponding to the observation of buckling in the plates, and to estimate the elastic stiffness under compression prior to this buckling event.

The main objectives were to find the load value corresponding to the observation of buckling in the plates, and to estimate the elastic stiffness under compression prior to this buckling event. As a result of this, the test was stopped after compressing 20 mm of the plate, which was found to be adequate to cause buckling in all the plates tested.(a)(b)Figure 8 (a) The servohydraulic Instron 8801 machine (50kN capacity), (b) Typical test specimen used in this study, tested under compression between platensTo make comparisons across many specimen designs, the buckling load and the initial, elastic plate stiffness was estimated. CAD file outputs were used to estimate the volume of each plate, which by multiplying by the published density of Nylon 12, yields a mass quantity.

CAD file outputs were used to estimate the volume of each plate, which by multiplying by the published density of Nylon 12, yields a mass quantity. This value was used to estimate a specific stiffness and a specific buckling load for each of the plates (the specific stiffness and specific buckling load are calculated by dividing each respective value by the calculated specimen mass). The buckling load was determined by finding the maximum force value from the response graph of each plate. Going backward from the buckling load to the start of the test, the stiffness of the plate was determined by finding the slope of the force- displacement curve. A linear regression curve was fit to the data region.

A linear regression curve was fit to the data region. If R2 was less than 0.99, the data range was reduced until this criteria was met, as shown in Figure 90.800.600.400.200.000.00y = 277.23x - 9.7339 R² = 0.994650.00100.00150.00200.00Figure 9 Example of a linear fit with R2 greater than 0.991899ResultsIn this section, load-displacement results from compression testing are presented for each of the five families of plates studied: baseline, triangular, square, stochastic, and topologically optimized. Each load-displacement graph, with the exception of the baseline, is accompanied with an image from a video that captured the deformation pattern under compression. In the next section, these datasets are normalized for comparison against each other and inferences drawn. Without normalization, the raw load-displacement plots may be misunderstood, since the mass of each of these plates is different.a.

Without normalization, the raw load-displacement plots may be misunderstood, since the mass of each of these plates is different.a. Unstiffened baselineEach of the six plates, separated in thickness by 0.5mm, shows an increase in peak buckling load, and stiffness as well, as expected, and shown in Figure 10. For all subsequent test results, the thinnest (1.5mm) and thickest (6mm) plates are represented as lower and upper bounds, respectively. This also helps compare all the graphs in relation to one another.Displacement (mm)2018161412108642002004006001.Smm Solid Plate2mm Solid Plate2.Smm Solid Plate8001000'§; :::.. "C ca 0 ...J3mm Solid Plate12004mm Solid Plate14006mm Solid Plate16001800Figure 10. Load displacement graphs for the six solid plate baselines (unstiffened) testedb.

"C ca 0 ...J3mm Solid Plate12004mm Solid Plate14006mm Solid Plate16001800Figure 10. Load displacement graphs for the six solid plate baselines (unstiffened) testedb. TriangularLoad displacement results for the five plates stiffened with the triangular patterns (two periodic, two graded, are shown in Figure 11a). Figure 11b shows the post-buckling profiles of the plates. As can be seen, all but pattern 5 have buckling initiate along the centerline. For pattern 5, the buckling is initiated above the denser graded region in the middle.c. SquareLoad displacement results for the three plates stiffened with the square patterns (one periodic, two graded, are shown in Figure 12a). Figure 12b shows the buckling profile of these plates. While the periodic square stiffened plate has a symmetric buckling pattern, the graded patterns clearly show a shift in the profile away from the center.1900d.

While the periodic square stiffened plate has a symmetric buckling pattern, the graded patterns clearly show a shift in the profile away from the center.1900d. StochasticThe stochastic shapes all showed fairly symmetric buckling shapes, as shown in Figure 13b. Pattern 9, with the highest localization of Voronoi cells in the center, had the highest buckling load, as seen in Figure 13a.e. Topology OptimizedThe topologically optimized plates showed results in line with expectations for the three mass retentions studied, see Figure 14a. Of all non-baseline shapes studied, the 50% mass retention optimized solution had the highest buckling load.

Of all non-baseline shapes studied, the 50% mass retention optimized solution had the highest buckling load. The next section will examine if this can be explained on the basis of the additional mass alone, or is improved because of design.Displacement (mm)20151050502501.5mm Solid Plate6mm Solid Plate450650;:::- C) .:,::.Periodic Triangular (Pattern 1)Periodic Triangular (Pattern 2)'ti -850 ns 0 .J1050Graded Triangular (Pattern 3)Graded Triangular (Pattern 4)1250Graded Triangular (Pattern 5)14501650(a)12345(b) Figure 11. (a) Load displacement graphs for the five plates stiffened with triangular patterns, along with the upper and lower baselines, and (b) Buckling patterns for the five patterns (side view for pattern 2 not available, hence showing front view during buckling)1901Displacement (mm)20181614121086420502501.5mm Solid Plate6mm Solid Plate450650850C Cl :::..

(a) Load displacement graphs for the five plates stiffened with stochastic patterns, along with the upper and lower baselines, and (b) Buckling patterns for the four patterns (side view for pattern 12 not available, hence showing front view during buckling)1903Displacement (mm)20181 614121086420502502mm Solid Plate6mm Solid Plate10% Mass Rel. TopOpt.450650 -Cl ::- "C n:, 0 ...J85030% Mass Rel. TopOpl.1 05050% Mass Rel. TopOpt.1 2501 4501 650(a)131415(b)Figure 14.

TopOpt.450650 -Cl ::- "C n:, 0 ...J85030% Mass Rel. TopOpl.1 05050% Mass Rel. TopOpt.1 2501 4501 650(a)131415(b)Figure 14. (a) Load displacement graphs for the five plates stiffened with topologically optimized patterns, along with the upper and lower baselines, and (b) Buckling patterns for the three mass retentionvalues (side view for pattern 2 not available, hence showing front view during buckling)1904DiscussionAs discussed before, load-displacement data presented in Figures 9-13 do not lend themselves to meaningful comparison since the mass of each plate varies. If one is to examine the effectiveness of any particular design strategy, a “specific” metric, that is normalized with respect to mass, may be used. The two metrics of interest in this work are the specific stiffness and the specific buckling load and these are plotted in Figures 15 and 16, respectively, color coded by the type of stiffening strategy employed.

Specific stiffness was obtained by dividing the estimated stiffness for each plate by the mass of the plate (estimated using the volume of the plate in design and the published density of Nylon 12 in the supplier datasheet). A similar procedure was followed for estimated buckling load, dividing the peak load by the mass of the plate.For specific stiffness, the results suggest that improvements in plate stiffness under compression over the baseline stiff plates, are attainable. The baseline plates show increasing stiffness with more mass, as would be expected, with the exception of the 6mm solid plate. As shown in Figure 14, using the highest performing baseline plate (4mm) as a reference (shown as the dotted line), several shapes outperform this baseline. The top 3 shapes that do so are labeled. For triangular shapes, the diagonal overlay (Pattern 2 in Figures 4 and 10) performs better than the cross-pattern (Pattern 1).

The top 3 shapes that do so are labeled. For triangular shapes, the diagonal overlay (Pattern 2 in Figures 4 and 10) performs better than the cross-pattern (Pattern 1). However, the top performing shape is somewhat surprisingly, the stochastic shape. Stochastic geometries tend to be more commonly associated with high- compliance, energy absorbing structures [16,17]. The uniform stochastic stiffener plate (Pattern 12) is not particularly impressive, but when gradation is added to the shape, it does have the promise of increasing specific stiffness, at least based on these results.8070e E 60 d, ~l 50 ., "' "' 40 C :t: ~ 30 <.: ·o ., 20V1Type - • • • -So lid Plate Triangular square stochastic TopOptQ. Vl10Figure 15. Specific stiffness histogram by design, classified by type of stiffening strategyWith regard to buckling load, the results suggest that no cellular stiffening strategy outperformed the solid plate, as can be seen in Figure 16.

However, within all the stiffening designs considered, the triangular Pattern 2 again outperformed the rest. Introduction of gradation in the1905center seemed to not positively effect the buckling load for these structures. The topology optimized solution with 50% mass retention had a high specific buckling load.120TypeCl 100 ;;::. C) ~ -0 80 RJ 0 ...J C) .!: ::;. u ::, a::i ;,.; 40 u ·;:; Q.> c.. V'> 2060So lid Plate • Triangular Square - Stochastic • TopOpt -Figure 16. Specific buckling load histogram by design, classified by type of stiffening strategyWhile this dataset is very limited, both in terms of design space explored, and in terms of replicates to build confidence in it, it suggests some promising lines of future inquiry. The periodic, diagonal triangular shape seems to be the best performing stiffening design when both specific stiffness and specific buckling load are considered.

The periodic, diagonal triangular shape seems to be the best performing stiffening design when both specific stiffness and specific buckling load are considered. Interestingly, localizing mass at the center, as was done with the step- and graded designs, seemed to have no improving effect on the data, suggesting that the aperiodicity of these designs may have led to a buckling instability that the more uniform geometries took longer to initiate. Stochastic stiffening also showed some promise – and there are some examples in nature (like dragonfly wings [18]) that suggest stochastic designs may be useful for thin plate stiffening, at least in the context of the combined goals of fluid flow and stiffening, and where buckling loads are less of a concern in relation to maintaining plate shape.AcknowledgementsThis work was conducted as part of a research project in the MFG 598: Design for Additive Manufacturing course offered in the Spring of 2019 at the Arizona State University.

The authors wish to acknowledge support from the Maricopa County Industrial Development Authority (MCIDA) and the Fulton Schools of Engineering at Arizona State University, where these parts were fabricated and tested, as well as Altair Engineering and nTopology for providing access to their software that was used to generate the designs in this paper.1906References[1] [2]D. Bhate, Four Questions in Cellular Material Design, MDPI Mater. (2019) 1–12. R. Chacón, E. Mirambell, E. Real, Transversally and longitudinally stiffened steel plate girders subjected to patch loading, Thin-Walled Struct. (2019). doi:10.1016/j.tws.2019.02.009. Z. Xu, G. Tong, L. Zhang, Design of horizontal stiffeners for stiffened steel plate walls in compression, Thin-Walled Struct. 132 (2018) 385–397. doi:10.1016/j.tws.2018.09.003. Q.-V. Vu, G. Papazafeiropoulos, C. Graciano, S.-E.

PERIMENTAL STUDY OF POLYMER ELECTROLYTE MEMBRANE FUEL CELLS USING A GRAPHITE COMPOSITE BIPOLAR PLATE FABRICATED BY SELECTIVE LASER SINTERINGNannan Guo and Ming C. Leu Department of Mechanical and Aerospace Engineering Missouri University of Science and Technology, Rolla, MO 65409, USAAbstractSelective Laser Sintering (SLS) can be used to fabricate graphite composite bipolar plates with complex flow fields for Polymer Electrolyte Membrane (PEM) fuel cells. The additive manufacturing process can significantly reduce the time and cost associated with the research and development of bipolar plates as compared to other fabrication methods such as compression molding. In this study, bipolar plates with three different designs, i.e., parallel in series, interdigitated, and bio-inspired, were fabricated using the SLS process. The performance of these SLS-fabricated bipolar plates was studied experimentally within a fuel cell assembly under various operating conditions.

The performance of these SLS-fabricated bipolar plates was studied experimentally within a fuel cell assembly under various operating conditions. The effect of temperature, relative humidity, and pressure on fuel cell performance was investigated. In the tests conducted for this study, the best fuel cell performance was achieved with a temperature of 75 (cid:1), relative humidity of 100%, and back pressure of 2 atm.1. IntroductionA fuel cell is an electrochemical cell that converts chemical energy from a fuel and an oxidant into electrical energy. A typical fuel cell system can operate at 40-60% efficiency, while a fossil fuel power plant operates at only about 35% efficiency. Additionally, a fuel cell system can be made very compact and lightweight without moving parts. These unique advantages of a fuel cell provide it with many application opportunities.

Additionally, a fuel cell system can be made very compact and lightweight without moving parts. These unique advantages of a fuel cell provide it with many application opportunities. Several types of fuel cells exist, including the polymer electrolyte membrane (PEM) fuel cell, solid oxide fuel cell (SOFC), molten carbonate fuel cell (MCFC), direct methanol fuel cell, and so on. Compared to other types of fuel cells, the PEM fuel cell has great advantages including low-temperature operation, high-power density, fast start-up, system robustness, and low emissions.A bipolar plate is a vital component in a PEM fuel cell. It supplies hydrogen and oxygen to reaction sites, removes the reaction product (i.e., liquid water) out of the fuel cell, collects the generated current, and provides mechanical support for the fuel cell. Bipolar plates constitute more than 60% of the weight and 30% of the total cost in a fuel cell stack [1].

Bipolar plates constitute more than 60% of the weight and 30% of the total cost in a fuel cell stack [1]. The bipolar plate significantly influences the performance of a PEM fuel cell. For example, the electrical conductivity of the bipolar plate itself affects the fuel cell’s internal resistance. The flow field design of the bipolar plate, which is used to transport reactants, greatly affects transport efficiency and, hence, fuel cell performance. Previous studies [2] have shown that the power density of a fuel cell can be increased significantly if the bipolar plate has a proper flow field design.212Bipolar plates usually are made of metal, graphite, and graphite composite. Metal has high electrical conductivity, but its poor chemical resistance does not allow it to persevere long within the chemical environment of a PEM fuel cell. Graphite has good electrical conductivity and chemical resistance, but its brittleness makes it very difficult to machine.

Graphite has good electrical conductivity and chemical resistance, but its brittleness makes it very difficult to machine. Given the ease with which they can be manufactured and their excellent chemical resistance, graphite/polymer composites are becoming increasingly more promising for producing bipolar plates [3-10]. Two fabrication methods for massively producing these bipolar plates are injection molding (IM) [3] and compression molding (CM) [4]. Although these methods are suitable for mass production, they are not time-efficient or cost-effective for studying the bipolar-plate flow field design, which includes channel cross-section shapes, dimensions and layouts. In the IM and CM processes, the mold corresponding to each different design has to be fabricated, which is very expensive and time-consuming.Selective Laser Sintering (SLS), as an additive manufacturing technology, has been studied to fabricate graphite composite bipolar plates for fuel cells [11-15].

Compared with CM and IM, SLS can fabricate graphite composite bipolar plates at a lower cost and shorter lead time, especially for the purpose of researching and developing new bipolar plate designs. In the SLS process, a mixture of graphite materials and binder in a powder bed is scanned by a laser, and the molten binder bonds graphite particles together to form 3D parts layer by layer. After a green part is built, it undergoes post-processing, including carbonization and infiltration, to increase the bipolar plate’s electrical conductivity, mechanical strength and gas impermeability. Studies have been conducted to investigate the properties of the bipolar plates fabricated using different graphite materials [16].

Studies have been conducted to investigate the properties of the bipolar plates fabricated using different graphite materials [16]. However, the performance of fuel cells with SLS-fabricated bipolar plates has not been investigated, nor has the influence of different operating conditions on the performance of these fuel cells that use SLS-fabricated bipolar plates.In this study, graphite composite bipolar plates with various flow fields, including parallel in series, interdigitated and bio-inspired designs, were fabricated using the SLS process. These fabricated bipolar plates were assembled in a PEM fuel cell unit. The performance of the fuel cells with these bipolar plates was studied experimentally under various operating conditions. The effects of temperature, relative humidity and pressure on fuel cell performance were investigated. The performance characteristics of fuel cells with different bipolar plate designs were compared.

The performance characteristics of fuel cells with different bipolar plate designs were compared. The long-term performance of the fuel cell using SLS-fabricated bipolar plates also was studied.2. Experiments2.1 MaterialsNatural graphite (3610) and synthetic graphite (4437), obtained from Asbury Graphite Mills, Inc. (New Jersey, USA), were used to fabricate graphite composite bipolar plates. Natural graphite has high electrical conductivity but poor flowability due to its flaky shape. Synthetic graphite is a product made by treating amorphous carbon materials, usually calcined petroleum coke and coal tar pitch, at high temperatures. Because of its more uniform and spherical shape, synthetic graphite provides better flowability in a packed powder. The material composition in this study was 55vol% NG, 10vol% SG and 35vol% binder. Phenolic powder (GP-5546, Georgia213Pacific) with an average size of 15µm was used as the binder.

The material composition in this study was 55vol% NG, 10vol% SG and 35vol% binder. Phenolic powder (GP-5546, Georgia213Pacific) with an average size of 15µm was used as the binder. The electrical conductivity was measured around 120 S/cm and the flexural strength around 40 MPa from a previous study [16]. After carbonization, porous brown parts were infiltrated with liquid epoxy resin (EPONTM Resin 828) in order to obtain gas impermeable parts.2.2 Fabrication processGraphite powder and phenolic binder were ball milled together for 12 hours to achieve a uniform mixture. An SLS machine (Sinterstation 2000) was used to build green parts from this mixture. The SLS process is depicted in Fig. 1. The CAD model of a bipolar plate was first converted into slices, and then the laser beam scanned the powder bed according to the cross- section of each slice (see Fig. 1). The binder, which has a low melting point, was melted by laser and bound the graphite particles.

1). The binder, which has a low melting point, was melted by laser and bound the graphite particles. After one layer was scanned, the powder bed was lowered by one layer of thickness, and a new layer of the powder material was fed on the top. The process was repeated until the bipolar plate was fabricated completely. The parameters used in the SLS process were as follows: fill laser power (14 W), outline laser power (4 W), laser scan speed (1.5 m/s), layer thickness (0.1 mm), and laser scan spacing (0.08 mm). The temperatures of the part bed and feed bins were maintained at 60 oC and 40 oC, respectively. For carbonization, green parts were heated to 1000 oC in a furnace filled with Argon gas to dissociate the binder and convert it to carbon residue to get brown parts. The heating schedule was from room temperature to 200 oC with a heating ramp rate of 60 oC/h, followed by a slower rate of 30 oC/h to 600 oC, and then a 50 oC/h ramp rate to 1000 oC, maintained for 1 h.

Brown parts, which are porous and weak, must be infiltrated with a liquid resin to become gas impermeable and to increase their mechanical strength. For infiltration, porous brown parts were immersed into the mixture of liquid epoxy resin and a curing agent, taken out after 20 min for their surfaces to be cleaned, and then placed in an oven at 80 oC for 30 min to cure the infiltrated resin.Figure 1. Fabrication process of Selective Laser Sintering.2.3 Bipolar plate designsThree flow patterns (parallel in series, interdigitated and bio-inspired designs) were developed in this study. The parallel in series design (Figure 2(a)) consisted of four sets of connected parallel channels with a channel width of 1.5 mm and a land width (distance between two adjacent channels) of 1.0 mm. The interdigitated design (Figure 2(b)) was composed of two214other to the outlet. sets of parallel channels, one of connected.

The interdigitated design (Figure 2(b)) was composed of two214other to the outlet. sets of parallel channels, one of connected. In this design, These two sets of channels were under the land area, which reactants are forced to flow through excessive pressure drop increases the reactant concentration on the catalyst surface in this design. The is required. The channel width was of leaves, which transport bio-inspired design (Figure 2(c)) (Figure 2(c)) was inspired by the vein structures of leaves, mirroring the function of mass efficiently from one central source to the whole surface of a leaf, mirroring mass efficiently from one central source to the whole surface of inspired design with three leaf vein structure was mimicked in bio-inspired design bipolar plate flow fields. The leaf vein structure ndary and tertiary generations. The width of all of generations of channels: the primary, seco primary, secondary and tertiary generations.

The leaf vein structure ndary and tertiary generations. The width of all of generations of channels: the primary, seco primary, secondary and tertiary generations. The width of all 1.5 mm in order to achieve a fair comparison with the other two designs. At these channels was 1.5 mm in order to achieve a fair comparison with the other two designs. 1.5 mm in order to achieve a fair comparison with the other two designs. the land area and the remaining reactant is forced to flow through the the end of all of the channels, the remain channel depth of all three merge to the outlet. The active area, thickness the three bipolar designs were the same at 50×50 plates are shown in plate designs were fabricated using the SLS process. The fabricated bipolar plates plate designs were fabricated using the SLS process Figure 3.one of which was connected to the inlet, and the other tointerdigitated together but not directly connected.

In thislow through the gas diffusion layer (GDL) under the lthe reactant concentration on the catalyst surface; however, an excessivewas 1.5 mm and the land width was 1.0 mm in this designactive area, thickness of bipolar plates, and channel depth 50×50 mm2, 3.0 mm and 1.5 mm, respectively. All(a) Figure 2. Bipolar plates with different flow field designs: (a)(b)(c) parallel in series design; (c)Bipolar plates with different flow field designs: (a) parallel in series interdigitated design; (d) bio-inspired design. interdigitated design; (d) bio(a)(b)(c)Figure 3. Bipolar plates with different flow field designsBipolar plates with different flow field designs fabricated using the SLS process:SLS process: (a)parallel in series design; (b) interdigitatedinterdigitated design; (c) bio-inspired design. All three bipolar plates All three bipolar plateshave an active area of 50×50 mm2.

All three bipolar plates All three bipolar plateshave an active area of 50×50 mm2. have2.4 Fuel cell testingA single fuel cell unit was assembled using two fabricated bipolar plates, two end plates, , as shown in Figure 4. Two gold- (from FuelCellsEtc) used to collect the generated current. The MEA (from FuelC mg/cm2 platinum as the catalyst on both the anode and cathode sides anode and cathode sides,was assembled using two fabricated bipolar platesand one commercial membrane electrode assembly (MEA), as shown in Figure and one commercial membrane electrode assembly (MEA) plated copper plates were used to collect the generated current. comprises a loading of 4 mg/cm215Nafion® 115 as the membrane, and carbon cloth as the gas diffusion layer (GDL). Fuel cell performance was measured using a test station (Greenlight Innovation, G40) under different operating conditions in order to study the effects of temperature, humidity, and pressure.

Fuel cell performance was measured using a test station (Greenlight Innovation, G40) under different operating conditions in order to study the effects of temperature, humidity, and pressure. The temperature varied from 55 (cid:1) to 85 (cid:1), the relative humidity for both the anode and cathode varied from 0 to 100%, and the back pressure varied from 0 to 2 atm. Table 1 provides details about the experimental designs. A thermocouple mounted inside the end plate was used to measure the temperature of the fuel cell. The flow rates of the anode and cathode were controlled by two separate mass flow controllers and were 0.3 Lpm and 1.05 Lpm, respectively, for all of the experiments. The relative humidity was controlled by adjusting the dew point temperature of the inlet flow. After the operating temperature and humidity were set, it took 90-120 minutes on average to reach the steady state.

After the operating temperature and humidity were set, it took 90-120 minutes on average to reach the steady state. After that, testing began by increasing the current density at each step and recording the corresponding voltage after two minutes (making sure the output was steady). The increase in current density continued until the cell voltage dropped to 0.3 V. After that, the current density was set back to zero, and two more tests were performed. The final data for each condition was obtained by taking the average of the three tests. The power density was obtained by multiplying the current density and the corresponding voltage.(a)(b)Figure 4. (a) Major components in a PEM fuel cell; (b) Actual fuel cell assembly used in the experiments.Table 1.

(a) Major components in a PEM fuel cell; (b) Actual fuel cell assembly used in the experiments.Table 1. Operating conditions for fuel cell testing experiments.Group A (temperature)Group B (relative humidity)Group C (back pressure)Temperature ((cid:1)) 55 65 75 85 75 75 75 75 75 75 75Relative humidityAnode 100% 100% 100% 100% 25% 50% 75% 100% 100% 100% 100%Cathode 100% 100% 100% 100% 25% 50% 75% 100% 100% 100% 100%Back pressure (atm) 0 0 0 0 0 0 0 0 0 1 22163. Results and Discussion3.1 Effect of temperatureThe performance of a PEM fuel cell is characterized by its polarization curve, which is the curve of voltage versus current density. Figure 5 shows the polarization curve as well as the curve of power density versus current density under different temperatures for the three flow field designs.

Figure 5 shows the polarization curve as well as the curve of power density versus current density under different temperatures for the three flow field designs. The performance of the fuel cell shows a typical polarization curve, indicating that the graphite composite bipolar plates fabricated using the SLS process are satisfactory for application to PEM fuel cells. All of the results show that, initially, the power density continually increases as the current density increases, reaching the maximum value when the current density is around 900 mA/cm2. As shown in Figure 5, the performance of the fuel cell greatly increases when the temperature increases from 55 (cid:1) to 65 (cid:1). After reaching 65 (cid:1), the performance of the fuel cell remains almost the same from 65 (cid:1) to 75 (cid:1), and then decreases from 75 (cid:1) to 85 (cid:1). All three designs show this similar trend in terms of the influence of temperature.

All three designs show this similar trend in terms of the influence of temperature. The parallel in series design experiences a large increase from 55 (cid:1) to 65 (cid:1) and a significant drop from 75 (cid:1) to 85 (cid:1), but for the interdigitated and bio-inspired designs, the differences between the two ranges of temperatures are relatively small.

As Equation (2) reveals, because Δ(cid:28)̂ is negative for the PEM fuel cell reaction, the thermodynamic voltage will decrease as the temperature increases. An increase in temperature also will reduce activation losses (due to reaction kinetics) due to the increase in exchange current density as the temperature increases. This is because increasing the reaction temperature will increase the thermal energy available in the system, and all of the particles in the molecule will move and vibrate with increased intensity. This higher level of thermal activity increases the likelihood that a given reactant will possess sufficient energy to reach the activated state, thus increasing the rate of reaction and decreasing activation losses. This dynamic can be expressed in the Butler-Volmer equation (Equation (3)) and the definition of exchange current density (Equation (4)), i.e.

The Butler-Volmer equation states that the current produced by an electrochemical reaction increases exponentially with activation overvoltage, which is sacrificed to overcome the activation barrier associated with the electrochemical reaction. Therefore, if more current is drawn from the fuel cell, more activation overvoltage is lost. If the exchange is increased to draw the same amount of current density from the fuel cell, less current density (cid:31)(cid:19) activation overvoltage will be needed. As Equation (4) indicates, increasing the temperature exchange will increase the current density exponentially, thus reducing activation losses. Therefore, the effect of temperature on fuel cell performance is the compromise of these two opposite influences, i.e., negative on thermodynamic voltage and positive on activation losses.

Therefore, the effect of temperature on fuel cell performance is the compromise of these two opposite influences, i.e., negative on thermodynamic voltage and positive on activation losses. When the temperature increases, the positive influence on the reduction of activation losses initially is greater than the negative influence on the reduction of thermodynamic voltage. Therefore, the overall performance of the fuel cell increases. However, after the temperature exceeds 65 (cid:1), these two effects begin to merge, and the overall fuel cell performance barely changes from 65 (cid:1) to 75 (cid:1) . Finally, the negative effect will surpass the positive effect, inhibiting the performance of the fuel cell as the operating temperature increases.

Finally, the negative effect will surpass the positive effect, inhibiting the performance of the fuel cell as the operating temperature increases. The experimental results for these different flow field designs agree well with the theoretical analysis and show that the range of 65 (cid:1) to 75 (cid:1) is the optimal operating temperature for all of these designs.is the decay rate from reactants to products, and Δ;4(a)(b)(c)(d)218(e)(f)Figure 5. Effect of temperature on fuel cell performance (Group A). The relative humidtity was kept at100%, and back pressure at 0 atm: (a, b) parallel in series design, (c, d) interdigitated design and (e, f) bio-inspired design.3.2 Effect of relative humidityThe performance of fuel cells with graphite composite bipolar plates fabricated using the SLS process was investigated under different relative humidity, with the temperature kept at 75 (cid:1) and the back pressure at 0 atm.

The experimental results for fuel cells with different designs are shown in Figure 6. Figure 6(a, c, e) shows the polarization curves under different relative humidity, and Figure 6(b, d, f) shows the corresponding power density curves. As indicated, relative humidity has a significant effect on fuel cell performance. The power density increased greatly as humidity increased from 25% to 75%; the rate of this increase slowed slightly after reaching 75%. The power density increased approximately 4-8 times when the humidity increased from 25% to 100%. All of the performance charts of fuel cells using the three different flow field designs show this same trend.Relative humidity mainly affects the ionic resistance of the polymer electrolyte (Nafion) because Nafion has to be hydrated by liquid water in order to maintain high proton conductivity.

The relationship between the proton conductivity of Nafion and water content is as follows [19]:4 K(cid:19)K (cid:8)4(cid:7)(cid:26)L< (cid:5) (cid:25)0.005193C (cid:8) 0.00326(cid:26)exp I1268(cid:25) where < is the conductivity of Nafion, C is the water content, and (cid:18) is the temperature. According to Equation (5), the proton conductivity of Nafion increases linearly with increase in its water content C, which occurs when the operating humidity increases. As a result, the fuel cell performance increases greatly with an increase in humidity, as seen in Figure 6, because of the reduction of ohmic losses.

The219(5)(6)(7)experimental results from the PEM fuel cells using SLS-fabricated bipolar plates agree well with the theoretical analysis of the effect of relative humidity on fuel cell performance.(a)(b)(c)(d)(e) Figure 6. Effect of relative humidity on fuel cell performance (Group B). Temperature was maintained at 75 (cid:1) and back pressure at 0 atm: (a, b) parallel in series design, (c, d) interdigitated design and (e, f) bio-inspired design.(f)3.3 Effect of back pressureFigure 7 shows the effect of back pressure on the performance of fuel cells using the parallel in series design (a, b), interdigitated design (c, d), and bio-inspired design (e, f). The results show that the performance of the PEM fuel cell increases greatly as the back pressure increases from 0 atm to 2 atm.

The results show that the performance of the PEM fuel cell increases greatly as the back pressure increases from 0 atm to 2 atm. This occurs because the increase of back pressure will increase the reactant concentration at both the anode and cathode, according to the following equation:220! (cid:5)a0b c (cid:5)10(cid:7)(8)The reactant concentration will greatly increase the reaction rate (Equation 4), thus increasing fuel cell performance. Additionally, the high back pressure applied on the outlet of the flow field will force the reactants to flow into the porous gas diffusion layer (GDL), which also improves the mass transport towards the catalyst layer. Therefore, more reactants will take part in the electrochemical reaction, resulting in higher performance. The high back pressure also benefits the removal of water droplets formed within the GDL and keeps the fuel cell from flooding [20].

The high back pressure also benefits the removal of water droplets formed within the GDL and keeps the fuel cell from flooding [20]. Figure 7(a) shows that the mass transport loss becomes prominent when the current density rises above 600 mA/cm2 for zero back pressure in the parallel in series design, while at pressures of 1 atm and 2 atm, the mass transport loss shows no remarkable increase even when the current density is as high as 1200 mA/cm2. The power density curve also demonstrates this point. The power density began to decrease when the current density rose above 800 mA/cm2 at a back pressure of 0 atm, but it continuously increased at back pressures of 1 atm and 2 atm. Similar trends were observed on the effect of back pressure for the other two bipolar plate designs.(a)(b)(c)(d)221(e)(f)Figure 7. Effect of back pressure on fuel cell performance (Group C).

Effect of back pressure on fuel cell performance (Group C). Temperature was maintained at 75 (cid:1) and relative humidity at 100%: (a, b) parallel in series design, (c, d) interdigitated design and (e, f) bio-inspired design.3.4 Comparison of performance for different flow field designsFigure 8 shows the performance comparison of PEM fuel cells using different bipolar plate flow field designs. Polarization curves and power density curves are shown in Figure 8(a) and (b), respectively. The bio-inspired design performed best among the three flow field designs, with a maximum power density of around 0.5 W/cm2. This is due to the higher mass transport capability of the bio-inspired design, which can be seen from the polarization curves that indicate less mass transport loss associated with this design. The maximum power density of the interdigitated design was 0.43 W/cm2, which was slightly higher than that of the parallel in series design.

The maximum power density of the interdigitated design was 0.43 W/cm2, which was slightly higher than that of the parallel in series design. The polarization curves show that the parallel in series design initially performed very similarly to the interdigitated and bio-inspired designs; at the middle region, it even exhibited slightly higher performance than the other designs. These findings indicate that the parallel in series design had a relatively high contact electrical conductivity between the bipolar plates and GDLs. However, after the current density increased above 800 mA/cm2, the parallel in series design suffered from a larger mass transport loss due to insufficient mass supply, while the performance of the interdigitated and bio-inspired designs did not drop sharply because of their better mass transport ability. Eventually, the bio-inspired design performed better than the other two designs, indicating that it has the most efficient transport ability.(b) Figure 8.

Eventually, the bio-inspired design performed better than the other two designs, indicating that it has the most efficient transport ability.(b) Figure 8. Comparison of the performance of fuel cells with different flow field designs (parallel in series, interdigitated and bio-inspired) under the same operating conditions (temperature of 75 (cid:1), relative humidity of 100% and back pressure of 0 atm).(a)2223.5 Fuel cell long-term performance testThe long-term performance of PEM fuel cells using SLS fabricated bipolar plates was also investigated to study the long-term work capability of the fabricated graphite composite bipolar plates. The parallel in series design was used in this study, with the operating conditions set to 75 (cid:1), 100% relative humidity, and zero back pressure. After steady operating conditions had been reached, the performance of the fuel cell was recorded every 10 seconds for 6 hours.

After steady operating conditions had been reached, the performance of the fuel cell was recorded every 10 seconds for 6 hours. The result is shown in Figure 9, in which the x axis is the operating time and the y axis is the variation of the cell voltage, while a constant current density of 750 mA/cm2 is drawn out from the fuel cell by the test station. This current density is the point at which the fuel cell with the parallel in series design can produce the maximum power density, as shown in Figure 8(b). From the result shown in Figure 9(a), the voltage continuously increased over the first 120 minutes from 0.515 V to 0.535 V; after that, the cell voltage remained relatively steady at 0.535 V, with very little variation, around d0.005 V. The fuel cell achieved its steady state around 120 minutes after the pre-set operating conditions had been reached.

The fuel cell achieved its steady state around 120 minutes after the pre-set operating conditions had been reached. This is why, in all of the experiments, the tests were started 120 minutes after the operating conditions were reached, as described in Section 2.3. Figure 9(b) details the performance from the 170th minute to the 179th minute. The voltage decreased suddenly around the 174th minute; 80 ms later, it returned to the steady-state voltage; then, around the 176th minute, another small drop occurred. These voltage drops were caused by the accumulation of liquid water in the flow channels. When the liquid water was removed from the fuel cell by the flow of the reactants, which can be observed from the transparent outlet pipe, the voltage returned to the steady state. Then, once again, the liquid water slowly accumulated inside the flow field, thus decreasing the voltage, and suddenly the voltage increased to the normal value as the water was dragged out of the channels.

This result indicates that the SLS-fabricated bipolar plates could provide a fuel cell with a steady, long-term performance.(a)(b)Figure 9. (a) Six-hour performance of the PEM fuel cell with SLS-fabricated graphite composite bipolar plates; (b) a detailed performance plot from the 170th minute to the 179th minute.4. ConclusionsGraphite composite bipolar plates were fabricated using the Selective Laser Sintering process, which can make flow fields with complicated geometries. The fabricated bipolar plates were assembled into a fuel cell unit whose performance was tested under different operating conditions, with the temperature varying from 55 (cid:1) to 85 (cid:1), relative humidity from 25% to223100%, and back pressure from 0 atm to 2 atm. The effects of temperature, relative humidity, and back pressure on the performance of PEM fuel cells were investigated.

The effects of temperature, relative humidity, and back pressure on the performance of PEM fuel cells were investigated. Fuel cell performance increased initially when the temperature increased from 55 (cid:1) to 75 (cid:1) , but after that, the performance decreased with increases in temperature. Both relative humidity and back pressure had a significant effect on the performance of fuel cells; the higher the humidity or back pressure, the higher the power density of the fuel cell. An analytical model based on PEM fuel cell thermodynamics and electrochemistry was used to help explain these experimental results. The performance characteristics of fuel cells using the parallel in series, interdigitated, and bio- inspired bipolar designs also were compared against each other, and the results show that the bio- inspired design performed best.

The long-term operation of the PEM fuel cell with the SLS fabricated bipolar plates produced a very stable output, indicating that there was no water flooding within the flow field.AcknowledgementThis project is funded by the National Science Foundation grant #CMMI-1131659.References[1] Tsuchiya, H., Kobayashi, O. Mass production cost of PEM fuel cell by learning curve. Int. J.Hydrogen Energy, 2004, 29: 985-990.[2] Kloess, J. P., Wang, X., Liu, J., Shi, Z., Guessous, L. Investigation of bio-inspired flow channel designs for bipolar plates in proton exchange membrane fuel cells. Journal of Power Sources, 2009, 188: 132-140.[3] Muller, A., Kauranen, P., Ganski, A., Hell, B. Injection moulding of graphite compositebipolar plates. Journal of Power Sources, 2006(154): 467-471.[4] Dhakate, S. R., Mathur, R. B., Kakati, B. K., Dhami, T. L. Properties of graphite-composite bipolar plate prepared by compression molding technique for PEM fuel cell.

perimental Study of Selective Laser Sintering of Parmax®Jagan R Rajan, Kristin L Wood Laboratory of Freeform Fabrication The University of Texas at Austin Austin, TX - 78712Nick Malkovich Research Engineer Mississippi Polymers Technologies Inc. Bay St. Louis, MS - 39520-9078AbstractSelective Laser Sintering (SLS) has been used to successfully process crystalline polymers, ceramics and metals. However, wide range of materials available and their applications have acted as an impetus to expand the applications of SLS to different material systems. This paper presents one such experimental study undertaken to explore the application of SLS to process Parmax® polymers – a family of high performance amorphous poly p- phenylenes. As a part of the collaborative study, various process parameters and material formulations were tried out to ascertain the feasibility of the process and the initial results obtained look promising.

As a part of the collaborative study, various process parameters and material formulations were tried out to ascertain the feasibility of the process and the initial results obtained look promising. This paper lists the material formulations tested and process parameters controlled. Also, analysis of the results of the experimental study and the outline of the next phase of research to be undertaken have been described. This study was motivated by the numerous applications of Parmax® in the electronics, defense and aerospace industries as well as the goal to expand the applications and utility of SLS.IntroductionRapid prototyping has become an essential tool in dealing with dramatically shortened manufacturing product life cycles. It replaces machine carved prototype models and patterns with machine-built components produced in a few hours, suitable for testing of form, fit and function.

It replaces machine carved prototype models and patterns with machine-built components produced in a few hours, suitable for testing of form, fit and function. Functional testing often places the rapid prototyping part in an operating assembly, requiring high strength materials. Unfortunately, current polymeric rapid prototyping materials such as nylon and polycarbonate, do not offer sufficient mechanical properties for high strength components, restricting functional testing up to this point and thus limiting the overall utility of rapid prototyping.To overcome this limitation, new materials with higher strength need to processed through Selective Laser Sintering(SLS). One such material, developed at MPT is its novel, high strength class of rigid-rod Parmax® Self Reinforced Polymer (SRP) resins.RRRRRRRRRRFigure 1.

One such material, developed at MPT is its novel, high strength class of rigid-rod Parmax® Self Reinforced Polymer (SRP) resins.RRRRRRRRRRFigure 1. Generalized Structure of Rigid-Rod Parmax® SRPs242The extraordinary hardness, stiffness and strength of Parmax® SRPs make them attractive for a wide range of uses that include molded articles and extruded sheet, stock and profiles for aircraft interiors, electrical and electronic components, surgical instruments, bushings, gears etc.; films and coatings for fiberless printed wiring boards, abrasion resistant coatings etc.; foams for high strength, low moisture absorption core and structural components and in blends and additives in which Parmax® SRPs are used to improve the performance of other polymers.

With its application extending into so many different areas, it becomes important to manufacture functional prototypes quickly and cheaply in order to demonstrate its capability and also arrive at optimized material formulations for each application. Thus being able to make Parmax® SRPs compatible to SLS is important from this point of view and this study aims at achieving the same.The primary objectives of this research effort are to (1) develop sinterable powder formulations based on MPT’s high-performance Parmax® SRP resins, (2) determine initial processing conditions for these powders via the Selective Laser Sintering (SLS) rapid- prototyping process, and (3) fabricate initial test coupons to demonstrate the high-integrity components that can result from this combination of material and processing.

However, the ultimate goal of the research is to come up with optimized powder and process parameters to produce components with unprecendented mechanical properties, thus paving the way for the commercialization of Parmax® SRP-based SLS formulations.The selection and development of powder formulations based on Parmax® SRP resins suitable for SLS processing were carried out at MPT under the aegis of an SBIR grant while determination of initial processing conditions and production of test coupons was conducted at the University of Texas.The paper is organized in the following way. First, a brief description of mechanical and chemical properties of relevant Parmax® SRP resins developed, current resin processing methods as well as the Selective Laser Sintering process are stated. Following this, the process parameters considered, material formulations tested and the initial experimental trials are presented. The results of the experiments are then analyzed.

Following this, the process parameters considered, material formulations tested and the initial experimental trials are presented. The results of the experiments are then analyzed. The paper concludes with a section on the contributions of this effort and an outline for the next phase of the research.BackgroundPowder Properties:MPT’s Parmax® SRPs, a new family of rigid-rod self-reinforcing polyphenylenes, possess a unique combination of high performance mechanical properties along with good processing capabilities. Research largely inspired and supported by the Air Force Materials Laboratory in the 1970’s and 1980’s has demonstrated that rigid-rod polymers possess exceptional mechanical properties[1], but most are either intractable or are soluble only in strong acid solvents.

Parmax® SRPs have been designed to overcome the processing deficiencies of other rigid-rod polymers via the use of carefully chosen pendant side chains to impart solubility and thermal processability to the normally intractable rigid-rod polyparaphenylene backbone. Thus, Parmax® materials can be dissolved in a variety of common organic solvents (methylene chloride, N-methylpyrrolidinone, phenyl ethers and esters, etc.) and, perhaps most importantly, can be thermally fabricated (compression molded, extruded, injection molded, etc.).Parmax® SRPs are prepared by proprietary processes that assure the isomeric integrity of the structure. A variety of substituent groups (R) can be utilized, so Parmax® SRP resins comprise a243versatile class of polymers, like polyimides, polyesters, and polyamides, within which a wide range of structural variation is possible.

We have surveyed many different derivatives including the thermoplastic rigid-rod Parmax®-1000 (benzoyl side chain structure) and Parmax®-2000 (4- phenoxybenzoyl side chain) homopolymers (Figure 2.) and the "tougher" and more processible Parmax®-1200 series of "kinked" copolymers. Currently it appears that the Parmax®-1000 and Parmax®-1200 derivatives will provide the first commercial Parmax® molding resin products.OOOOn Parmax®-1000n Parmax®-1200x Parmax®-2000y nFigure 2. Chemical Structures of Some Thermoplastic Parmax® SRP DerivativesParmax® SRP resins demonstrate extraordinary strength, modulus, hardness, and other properties in a combination unrivaled by any other single material (Table 1). Metals are widely used for structural support but are heavy and subject to corrosion and fatigue. Common thermoplastics are inexpensive to fabricate into parts but currently cannot provide enough strength or stiffness for many applications.

Common thermoplastics are inexpensive to fabricate into parts but currently cannot provide enough strength or stiffness for many applications. Fiber-reinforced composites have evolved to offer unprecedented strength-to-weight characteristics but are extremely difficult and expensive to process. As demonstrated in Figure 3, isotropic Parmax® resins possess unparalleled mechanical strength and modulus when compared to other engineering thermoplastics. In fact, on weight basis, Parmax® polymers are extremely competitive with structural metals such as aerospace grades of aluminum and titanium (Figure 4).Table 1.

Mechanical Properties of Engineering Plastics and Parmax® SRPsk (h t g n e r t S3020d l e i Y100SteelAl-5052 Al-6061 Al-2024 Ti-6Al-4VParmax“Figure 4. Specific Strength Comparison of Various Metals with Parmax® SRPsProcessing Techniques:Ultra high performance, isotropic Parmax® SRP polymers are processed utilizing conventional polymer techniques. These include routine extrusion, thermoforming and compression molding operations. Injection molding grades are under development and their availability is anticipated for fourth quarter of 2001. Processing temperatures are in the range of 290 to 340 °C. Within this processing window, Parmax® SRP resins display remarkable stability and do not require any additional stabilizers to facilitate the processing.

Within this processing window, Parmax® SRP resins display remarkable stability and do not require any additional stabilizers to facilitate the processing. The excellent solubility of these materials in common organic solvents allow for the facile preparation of films and coatings from solution.Selective Laser Sintering Process:Selective Laser Sintering is a rapid prototyping process in which powdered material is melted into the desired shape through the repeated scanning of cross-sectional areas that will eventually form the 3D model[14].

The machine (figure 5) consists of one part piston which translates within the part cylinder; 1 or 2 feed pistons that move within the powder cartridge and act as the powder feeding/collection system; a roller mechanism for spreading powder; heating coils, sensors and controllers for heating and maintaining the powder at the correct temperature;245laser and its accompanying optics to focus the laser beam precisely; blower, sensors and controllers that maintain the process chamber inert(not shown in figure); a computer interface that serves as the communication link between the user and the machine(not shown in figure). The SLS process is accomplished with a CO2 laser which scans a cross section of the part on the surface of the part bed. After one cross section is scanned, the part piston moves down and the feed pistons, carrying powder above them, move up.

After one cross section is scanned, the part piston moves down and the feed pistons, carrying powder above them, move up. Then a roller sweeps across the bed, laying a fresh layer of powder and the whole process is repeated till the whole component is made. Uneven or large temperature gradients need to minimized during the build so as to prevent excessive part curl. This is achieved by preheating the powder in the feed bins as well as the part piston to prescribed temperatures.Figure 5: Schematic of the Selective Laser Sintering ProcessApproachThe aim of the study, as stated before, is to process Parmax® SRPs using Selective Laser Sintering. There are many powder as well as process parameters that affect the quality of components produced. Hence, to manufacture functional prototypes, these parameters have to be optimized to achieve the desired values of performance variables(responses) such as strength, modulus and dimensional accuracy.

Since the SLS process has not been mathematically modeled to the extent desired, a set of experiments need to be performed to identify important parameters and establish mathematical relationships between them that will help optimize the process.The first step in the research process was to identify the parameters that could affect the process. An initial study revealed 12 SLS process parameters and 4 powder parameters that could be controlled. These are briefly discussed below. SLS Process ParametersThe 12 SLS parameters are powder age, powder type, part bed temperature, feed piston temperature, part piston temperature, laser power, powder layer thickness, feed rate, chamber atmosphere, part orientation, roller speed and time between layers.246The first parameter to consider in the process is the type of powder to be used. There are currently 12 commercially available powders that can be used with the SLS process.

There are currently 12 commercially available powders that can be used with the SLS process. Each of these have different physical properties and chemical properties and produce different results. This study focuses on only Parmax – a family of high performance amorphous poly p- polyphenes.The designer then has the choice of using either new powder or recycled powder. The powder in the machine is pre-heated very close to its melting or softening temperature causing partial melting. This might cause clumping and hence change the particle size. Therefore, every time the powder is used, different laser power and pre-heating temperatures might be needed. Temperatures of the part bed, part piston and feed pistons have been observed to have a great effect on the properties of the final part. Normally, the feed pistons and part pistons are kept at temperatures below the part bed temperatures.

Normally, the feed pistons and part pistons are kept at temperatures below the part bed temperatures. This is to minimize powder clumping that might occur due to prolonged exposure of powder very close to its melting or softening point. The feed piston temperature and the part bed temperature have an effect on part curl, strength of part and on the dimensional accuracy.The laser power is another important parameter. The laser power can be controlled by varying the incident laser power, the scan speed of the laser and the scan spacing of the laser. However the machine allows the user to change only the incident power. The laser power like the piston and part bed temperatures has a profound effect on the part curl, strength and dimensional accuracy.The orientation is an important factor in determining the quality of the part. Because the build proceeds in the z direction(upward), taller part take longer to build.

Because the build proceeds in the z direction(upward), taller part take longer to build. Hence the build time can be minimized by orienting the longer dimension of the part in the X or Y direction(parallel to the bed of the machine). Orientation may also effect the part strength because in the Z direction, there is a potential problem with proper bonding between the layers. This is also expected to cause a decrease in strength along the Z direction. Surface roughness is also effected by the part orientation. The layered nature in which parts are produces causes a “stairstep” effect on the surfaces that are not parallel or perpendicular to the part bed surface[14].Layer thickness is a controllable parameter that effects many part properties. Among these are build time, strength, density and accuracy. The layer thickness is controlled by the movement of the part piston. More the movement of the part piston downward, more is the layer thickness.

The layer thickness is controlled by the movement of the part piston. More the movement of the part piston downward, more is the layer thickness. Coupled with the layer thickness is the feed rate which is determined by the distance moved by the feed pistons upward. The convention used in the process to ensure a uniform layer to be spread across the part piston is to have a feed rate that is twice the layer thickness. However, the particle size may effect the feed rate setting that is needed.The roller speed and the time between layers are parameters that have an effect on the build time and is suspected to have an impact on the part curl. This is because the roller speed determines the time in which the powder is spread across the part piston and also the force it applies on the sintered part while the time between layers determines the amount of time the freshly laid powder has to attain the part bed temperature.

If enough time is not given for the freshly laid powder to attain the part bed temperature, large temperature gradients might occur leading to excessive part curl.An inert atmosphere in the chamber is required during processing to prevent oxidation that might cause loose powder to solidify. This is achieved by supplying a steady flow of nitrogen to the build chamber. The percentage of oxygen present in the chamber effects post processing and accuracy of the part. Since it is known that a more inert atmosphere would247always be preferred, a steady nitrogen flow is maintained to limit the oxygen content below 5.5%(default machine parameter value).Powder ParametersThe four important powder parameters are particle size, density, viscosity and chemicalstructure.Powder particle size determines the time taken for the powder to melt when sintered. Also it is related to the minimum powder layer thickness and the feed rate.

Also it is related to the minimum powder layer thickness and the feed rate. Large particle size causes an increase in the layer thickness and leads to poor consolidation and hence strength. Thus it is important to arrive at an optimum particle size.The density of the powder is known to have a direct impact on the strength of the processed component and also on the degree of consolidation. Typical densities of powders that have been processed are between 0.4 gm/cm3 to 0.6 gm/cm3Viscosity determines the ease with which the powder flows and it is desirable that the powder exhibits low viscosity at the processing temperatures. However, modification of viscosity often entails a change in the chemical structure.Chemical Structure of the powder is one of the most important parameters but often its modification is difficult and restricted.

The structure of the polymer determines its viscosity, crystallinity and melting temperature.Performing an exhaustive set of experiments taking all of the 16 variables into account is time consuming and a waste of resources. In order to identify those variables that have a major effect on the response, an initial set of experiments were performed. The work progressed on dual tracks. One was formulation of Parmax® SRP resins suitable for SLS and the other was SLS process optimization.Initial formulation selectionResearch effort executed under the Phase I SBIR Contract initiated with assessment of the critical parameters of the successful SLS formulations and the fit that current Parmax® SRP resin offer. The first parameter examined closely was the effect of the powder particle size. The focus was on the identification of Parmax® SRP resins with a range of particle sizes to bracket the acceptable limits on this parameter.

The focus was on the identification of Parmax® SRP resins with a range of particle sizes to bracket the acceptable limits on this parameter. The initial two formulations represented materials with mean particle sizes varying from 8 µm (1200-341) to 60 µm (1200-367). Bulk densities for the two formulations were 0.02 gm/cm3 for 1200-341 and 0.25 gm/cm3 for 1200-367 Micrographs of these two formulations can be found in Figures 6 and 7, respectively.Figure 6. 1200-341 powder structure and size248Figure 7. 1200-367 powder structure and sizeBoth of the formulations utilized in this first SLS iteration are the Parmax® 1200derivative, the chemical structure of which can be found in Figure 8.Figure 8. Structure of Parmax® 1200Parmax® 1200 is a copolymer with about 15% of the flexible meta-phenylene linkages present. This intentional kink introduces a degree of flexibility to the otherwise rigid para- phenylene backbone.

This intentional kink introduces a degree of flexibility to the otherwise rigid para- phenylene backbone. This limited flexibility is translated into improved processability and molecular flow above glass transition temperature (Tg), when compared to the completely rigid Parmax® 1000.Initially, three materials were planned for the first iteration, with the remaining formulation being Parmax® 1200-367 with its particle size further reduced. This was attempted through the use of jet milling at Fluid Energy in Hatfield, PA. This vendor, however, failed to induce any particle size reduction. This was particularly puzzling, since the same vendor previously ground Parmax® 1200 to the 10-20 µm particle size specification.

This vendor, however, failed to induce any particle size reduction. This was particularly puzzling, since the same vendor previously ground Parmax® 1200 to the 10-20 µm particle size specification. As a result of this failure, only 2 Parmax® SRP formulations were tested during the first iteration of this effort.Initial sintering trialsInitial experiments were conducted using Parmax® 1200-367 to ascertain the feasibility of the process, reduce the number of process parameters by eliminating variables that had little effect on consolidation and to get the outer bounds on the important parameters. These experiments were driven by intuition and the insights obtained during experimentation and not by following formal experimentation methodologies.

These experiments were driven by intuition and the insights obtained during experimentation and not by following formal experimentation methodologies. This was mainly because an exhaustive set of experiments could not be run with the powder available and the goals of the experiment did not warrant a large set of experiments to be run.DSC curves of Parmax® 1200 indicated the onset of softening temperature (Tg) as 152 °C. With knowledge gained from processing parts using DuraForm (nylon-based SLS material), the part heater temperature was initially set at 132 °C, the left and right feed pistons were at 60 °C and a laser power of 5.5 watts (default power setting for DuraForm) was used. With these settings the powder did not consolidate. Hence the laser power was increased to 6 watts (at z249(measure of part thickness) = 0.012”). This did not alter the result and subsequently (at z = 0.016) it was increased to 9 watts.

Hence the laser power was increased to 6 watts (at z249(measure of part thickness) = 0.012”). This did not alter the result and subsequently (at z = 0.016) it was increased to 9 watts. This sintered the powder to a greater extent but the sintered powder was turning brown (which at this point was thought undesirable). To minimize the discoloration the part heater temperature was increased to 132 °C (at z=0.02”) and various laser powers were tried out. Also, part heater temperature was ramped up by 2 °C and at z = -.524” it was increased further by 4 °C to 142 °C. However, these changes brought about little change in the result. One main reason for not increasing the part heater temperature appreciably was to avoid solidification of all the powder in the part piston leading to wastage of large amounts of expensive powder.

One main reason for not increasing the part heater temperature appreciably was to avoid solidification of all the powder in the part piston leading to wastage of large amounts of expensive powder. An interesting phenomenon that was noticed during the experiment was that at around 0.028”, the sintered powder started warping and at 0.056” it was high enough to lead to movement of the part under the roller. Warping could occur due to the following reasons:(1) localized expansion due to uneven temperature gradients while scanning by the laser. (2) Excessive force applied by the movement of roller over the sintered part (3) uneven expansion due to a large difference in temperatures between the feed pistons and part pistons.Green parts obtained in these first iterations can be seen in Figure 9. Following thisstage, the parts were sent to Bodycote IMT for an additional HIP-ing step.Figure 9.

Following thisstage, the parts were sent to Bodycote IMT for an additional HIP-ing step.Figure 9. First Parmax® SRP parts produced via SLSThrough this experiment it was established that the grain size of the 1200-367 powder formulation was suitable for the SLS process, at least for the initial levels of consolidation. The roller was able to spread the powder evenly and bulk density variability was minimized. Based on the results of this run, overall powder bulk density and viscosity of the powder at these temperatures might have to be varied to optimize the consolidation.

Based on the results of this run, overall powder bulk density and viscosity of the powder at these temperatures might have to be varied to optimize the consolidation. From the point of view of the process parameters, the operating temperatures could be increased, other parameters like roller speed, power layer thickness and feed distance need to be controlled so as to determine their impact on warping and powder consolidation.The second material formulated for this iteration, Parmax® 1200-341, could not be adequately processed owing to the extremely low bulk density of this powder.

Even though it appeared that the small particle size would be beneficial from the standpoint of better surface finish as well as larger particle surface area for laser to operate on, the flocculent nature of this material did not allow the Sinterstation 2000 roller to distribute the powder evenly.Formulations aimed at viscosity reductionBased on results obtained during the first half of the effort, a second iteration of materials was formulated for the laser sintering trials. The stated goal during this phase of the effort was an improvement of the molecular flow of Parmax® SRP resins at the Sinterstation operating temperatures. Formulations were as follows:250Plasticized Parmax® SRP formulation (1200-366), containing 15% of di-octyl phthalate (DOP), an esteric plasticizer commonly used in PVC industry. DOP and Parmax® SRP resin were compounded using a 1” Davis Standard Killion Single Screw Extruder (SSE). Approximately 6 lbs of this formulation was produced and pelletized.

DOP and Parmax® SRP resin were compounded using a 1” Davis Standard Killion Single Screw Extruder (SSE). Approximately 6 lbs of this formulation was produced and pelletized. Pellets were then sent to Particle Reduction Corp of Elk Grove Village, IL for another attempt at particle reduction. The two processes used were hammer milling and jet milling. Hammer milling successfully ground Parmax® SRP pellets to approximately 500 µm particle size (consistency of beach sand). Unfortunately, further milling, utilizing the jet milling facilities, was again unsuccessful.Formulation utilizing a different copolymer of Parmax® SRP materials, Parmax® 1215K (Formulation designation 1215-373), structure of which can be found in Figure 10. This chemical structure modification had a goal of reducing the viscosity of the resulting material at Sinterstation processing temperatures, and also to examine the suitability of chemistry modification as means for optimizing sinterable Parmax® SRP derivatives.

Viscosity reduction is achieved through an addition of benzoxy-substituted meta- phenylene linkage. This hinders efficient packing of the rigid rod backbone chains, resulting in increased disorder within the polymer lattice. Resulting free volume increase translates into lower viscosity at processing temperatures of interest. 1215-373 had particle size of approximately 70 µm with bulk density of 0.23 gm/cm3.Figure 10. Chemical structure of Parmax® 1215KSintering studies of 1200-366 and 1215-373As a baseline for this round of sintering trials, conditions that led to the consolidation of parts in the first iteration were replicated. It became immediately obvious that 1200-366’s sand- like powder consistency was unsuitable for laser sintering process, owing to the inability of the Sinterstation roller to distribute the material evenly across the sintering surface.

The second formulation, labeled 1215-373 was successfully consolidated using the same parameters utilized in previous successful consolidation, 1200-367. Following the baseline establishment, the remainder of the machine time was devoted to the sintering parameter optimization, utilizing 1200-367 and 1215-373 materials.251Effect of the powder layer thickness on part consolidation and warpingLocalized expansion arising from uneven temperature gradients while scanning by the laser might be affected by process parameters such as powder layer thickness, time between layers, feed distance, part heater temperature, laser power used and the left and right feed temperature. The third experiment concentrated on studying the effect of powder layer thickness, feed distance and part heater temperature.Left and right feed distances were set initially at 0.014” and the powder layer thickness was increased to 0.006”.

The left and right feed heaters were set at 120°C during the warm up stage and at 145°C during the sintering stage. However, owing to the technical difficulties with Sinterstation 2000, the powder could not be heated beyond 115°C. The part heater was set at 154°C and laser power used was 6 watts. The increase in the powder layer thickness and the feed distance did not seem to have an impact on the warping and so the powder layer thickness and feed distance were decreased to 0.005” and 0.012” respectively (at z=0.18”) while the part heater temperature was ramped up to 165°C to aid consolidation of powder. To prevent discoloration of the powder, the laser power was not increased beyond 7.5 watts.Based on these results, we suspect that powder layer thickness and feed distance were not the primary reasons for uneven temperature gradients and localized expansion leading to warpage.

Also, we conclude that the temperature reached by the powder on preheating to 165°C and application of laser power of 7.5 watts is not high enough to sinter it completely.Impact of roller motionThe next experiment focused on studying the effects of the roller motion over the sintered part and also the time elapsed between laying of fresh powder by the roller and the start of the sintering process. The movement of the roller carrying powder might apply a force on the sintered part that is enough to cause yielding. So in this experiment the roller speed, and the amount of powder it carries, which is directly related to the feed distance, are varied, as is the time between which layers of powder are laid down.During this run, the left and right feed distances were set at 0.012”; the feed heaters were at 120 °C while the part heater was set at 175 °C (10 degrees more than the previous experiment). The laser power was maintained at 6.5 watts through out the experiment.

The laser power was maintained at 6.5 watts through out the experiment. The roller speed was at the default value of 5 in/sec but the time between layers was set at 50 sec (default value 15sec). At various points in the experiment, the roller speed was varied from a minimum of 3 in/sec to 7 in/sec and the feed distance between 0.012” to 0.035”. The effect on warping seemed minimal. The part heater temperature was increased to 180 °C and then to 190 °C. At these high temperatures, the powder in the part piston had started clumping and hence the roller was unable to spread the powder evenly. The clumping might have occurred because the temperature was too high or because of prolonged heating (the powder was at these temperatures for around 3 hours and being amorphous in nature might have started softening and coagulating gradually).Laser power optimizationThrough this experiment we aimed at obtaining the outer bounds for the laser power.

The part heater, piston heater and feed heater temperatures were set at 165°C, 130°C and 100°C respectively. Laser power was varied from 6 watts to 30 watts. At values beyond 15 watts, the laser power was high enough to totally degenerate the powder. The powder was getting252completely charred at these high values, establishing the upper bound for the laser power application to Parmax® SRP resins.From the above experiments, we conclude that the most important parameters that seem to impact sintering as well as warping are the feed piston temperatures, part heater temperature and laser power. Overall, these experiments were successful in eliminating 3 parameters (time between layers, roller speed and piston heater temperature). Also, it helped us obtain the outer bounds of part heater temperature and laser power.

Also, it helped us obtain the outer bounds of part heater temperature and laser power. This information is going to be crucial in reducing the number of variables to be examined during the next phase of research, therefore speeding up the screening process and allowing us to concentrate on mission critical parameters.ResultsDuring the first phase of study, we have demonstrated successful processing of Parmax® SRP resins using Selective Laser Sintering (SLS) techniques, offering the prospect of rapid prototyping of components with unprecedented mechanical properties. This phase also elucidated the requirements for development of commercially practical Parmax® SRP SLS materials. The research performed thus far consisted of these distinct phases:Materials Formulation: The initial phase involved selection and development of sinterable powder formulations based on Parmax® SRP resins that would be suitable for processing by Selective Laser Sintering (SLS) rapid-prototyping process.

Four candidate powder formulations were screened for sintering. Decisions to focus on these four candidate formulations were based on a number of properties of interest, most importantly powder particle size, viscosity, chemical structure and bulk density.SLS Trials: Following the identification of the candidate formulations by MPT, initial processing studies at The University of Texas at Austin were performed. Following the initial parameter determination, a set of conditions were selected and the initial fabrication of the parts was undertaken. This formulation iteration consisted of 2 formulations, one of which failed due to the low bulk density and tackiness of the powder (Parmax® 1200-341). Using the second formulation, Parmax® 1200-367, with a bulk density an order of magnitude higher than the first attempt, initial test coupons were successfully fabricated. Both of these formulations are based on Parmax® 1200, structure of which is illustrated in Figure 8.

Both of these formulations are based on Parmax® 1200, structure of which is illustrated in Figure 8. Dimensions of the first parts were approximately 30 mm × 7 mm × 1-3 mm. Following a successful consolidation phase, green parts were subjected to Hot Isostatic Pressing (HIP) at Bodycote IMT of London, OH.Mechanical Properties Assessment: Following this densification step, initial mechanical properties were measured(by standard 3 point bend test) and are reported in Table 2, along with typical properties of nylon and polycarbonate prepared by the SLS process. This data confirms the hypothesis that the intrinsically higher modulus and strength properties of Parmax® SRP resins would translate into higher mechanical properties of the SLS processed components when compared to the current materials in use.

This is even more remarkable when considering that the bulk density of Parmax® SRP resin utilized in these experiments is less than 50% of the conventional SLS materials.253MaterialPowder Bulk DensitySLS PropertiesTensile StrengthFlexural ModulusNylonPolycarbonate Parmax®++++ 12000.5-0.6 gm/cm3 0.6 gm/cm3 0.25 gm/cm35,200 psi3,400 psi~10,000 psi0.13 Msi0.15 Msi0.40 MsiTable 2. Comparison of mechanical properties of Parmax® 1200 and other SLSmaterialsBased on the knowledge base acquired during the first effort iteration, a second set of formulations was produced in an attempt to improve on the sintering characteristics observed during the initial trials. These included a structurally modified component, Parmax® 1215-373 (Figure 10), and a plasticized formulation to reduce the viscosity (Parmax® SRP 1200-366).

These included a structurally modified component, Parmax® 1215-373 (Figure 10), and a plasticized formulation to reduce the viscosity (Parmax® SRP 1200-366). One of the formulations, Parmax® 1200-366, underwent particle reduction via a hammer mill in order to obtain the powder necessary to attempt laser sintering. The second iteration of the sintering trials succeeded in fabricating parts of comparable quality to 1200-367 using Parmax® 1215-373, while the plasticized Parmax® 1200-366 proved very difficult to feed, owing to the very coarse grain structure resulting from the particle reduction operation. Optimization of the laser sintering conditions was relatively limited owing to the small supply of candidate resins, as well as some machine limitations with Sinterstation 2000 at University of Texas.

However, preliminary sintering condition optimization successfully eliminated several variables, thus rendering the next phase optimization effort a much less time intensive task.Conclusion and Future WorkIn summary, this research effort was successful in demonstrating that Parmax® SRP resins can be consolidated utilizing the SLS process. Sintered components were tested for mechanical properties and, as expected, far exceed mechanical properties obtainable with conventional SLS polymeric materials, even in their current un-optimized state, thus fulfilling all three of the objectives. Work performed on a structurally modified Parmax® 1215-373 also opens up new and exciting possibilities for optimizing the sinterability of Parmax® SRP resin even further through judicious chemistry and structure manipulation. This reinforces the promise that these materials hold for Rapid Prototyping applications following the optimization during the next phase of research.

This reinforces the promise that these materials hold for Rapid Prototyping applications following the optimization during the next phase of research. Commercial potential of this development will be especially significant in laser sintering of a wide variety of low-volume, high integrity components for many possible commercial and industrial prototype needs, as well as prototyping needs in aerospace, military and tooling areas. The ultimate goal of this project is to produce materials strong enough to be utilized as fully functional prototypes, not only visual design aides, and thus fulfill one of the most vexing needs of the laser sintering community.

Parmax® SRP materials appear ready to be the materials the rapid prototyping community relies on to achieve these ambitious goals.+ Properties measured after Hot Isostatic Pressing254Recommendations for further work involves continuing focus on optimization of Parmax® SRP materials that were initially processed by laser sintering techniques during the current study as well as the optimization of the Laser Sintering processing parameters. Therefore, a dual track approach is recommended:Optimization of resin molecular architecture and composition (e.g., molecular weight and distribution, melt rheology, additives, co-monomer selection, degree of crystallinity etc.) and powder characteristics (e.g., particle size and distribution, bulk density, etc.) to enable effective sintering with high retention of mechanical properties.

This work will be performed under a rigorously designed experimental matrix in order to maximize the efficiency of the research effort.An optimal set of laser sintering conditions will be established. This will draw on the preparation of initial test coupons fabricated at the University of Texas at Austin during the Phase I effort, and will be performed on statistically based Design-Of-Experiments (DOE) principles. This work will be performed at University of Texas’ Laboratory for Free Forming (where the technique was invented) to insure successful Laser Sintering parameter establishment.References 1. Adams, W.W.; Eby, R.K.; McLemore, D.E.; Editors "The Materials Science and Engineering of Rigid-Rod Polymers," Mat. Res. Soc. Symp. Proc. 1989, Volume 134. 2. Beaman J. J.; Barlow J. W.; Bourell, D. L.; Crawford R. H.; Marcus, H. L.; McAlea, K. P.; Solid Freeform Fabrication: A new direction in Manufacturing: With research and applications in Thermal Laser Processing.

perimental Study of Snap-Fits Using Additive ManufacturingKevin Torossian1 and David Bourell21National Engineering School of Saint-Etienne, 42023 Loire, France 2Laboratory for Freeform Fabrication, Mechanical Engineering Department, The University of Texas at Austin, TX 78712AbstractA snap-fit is a mechanical joint system whose mating parts exert a cam action, flexing until one part slips past a raised lip on the other part, preventing their separation. The use of snaps in additive manufacturing (AM) is an approach for assembling components of parts too large to build in one piece in AM. There are broadly two types of snap-fits possible to encounter, permanent and non-permanent, depending on the design geometry. An experimental study was carried out to evaluate the mating/dismounting force for snap-fits regarding several geometrical parameters for additive manufacturing. The design chosen for this study has been established from the start to work on only one design.

The design chosen for this study has been established from the start to work on only one design. The parameters chosen for experimental investigation were the mating angle, the separation angle and the inner diameter of the mating part. All in all, fifteen pairs were designed and additive manufactured for evaluation. The force required to insert and separate the snap components was recorded and compared to the value based on a derived equation.IntroductionSnap-fits are a simple, quick and cost effective method of assembling two parts. When designed properly, parts with snap-fits can be assembled and disassembled numerous times without any effect on the assembly. Depending on the application, snap-fits can also create a permanent assembly. Snap-fits are also an environmentally friendly form of assembly because of their ease of disassembly, making components of different materials easy to recycle. Elimination of adhesives is also an environmental advantage.

Elimination of adhesives is also an environmental advantage. Although snap-fits can be designed with many materials, the ideal material is thermoplastic because of its high flexibility. Other advantages include its relatively high elongation, low coefficient of friction, and sufficient strength and rigidity to meet the requirements of most applications.There has always been a desire to build parts that are larger than the AM fabrication build chamber. Typically components are glued and/or dove tailed together after building. The purpose of this experimental study was to verify if snap-fits fabricated with AM could be used to attach components to one another, for plastic parts at least. Applications exist where the assembly would be permanent, or in some cases, it could be desirable to take the assembly apart.

Applications exist where the assembly would be permanent, or in some cases, it could be desirable to take the assembly apart. That is why the required force for snapping and unsnapping is important.The development of equations for the detailed sizing of snap-fit features and for predicting their response can be accomplished using analytical, numerical or experimental methods. As shown in Figures 1 and 2, several types of design for snap-fits exist including different cross sectional shape and thus diverse equations to calculate the mating/dismounting force [1].Figure 1: Different examples of snap-fits design [1]Figure 2: Equations regarding snap cross sectional shape and design [2]Figure 3 illustrate the snap-fit design chosen for this experimental study.

The cross section is a circular cross section constant over the length.Figure 3: Chosen design [3]Due to cost and time issues, the experimental method is usually replaced by a numerical method using the finite element method to simulate experiments. In this case, analytical and experimental results have been compared.

With this derived equation, the mating/dismounting force can be analytically calculated by altering geometrical parameters and thus positively impacting the snap-fit design. According to prior research [1], the equation to have the dismounting force, 𝐹𝑑 , is the same as the one above but with 𝛼 being replaced by 𝛼′.Moreover, the derived equation takes into account not only the geometrical parameters but also the mechanical properties of the material such as the Young’s modulus and the friction coefficient.ExperimentalAs mentioned earlier, only one design has been chosen, and, to see the most representative evolution of the force, some geometrical parameters have been set. On the other hand, the inner diameter 𝐷0, the mating angle 𝛼 and the separation angle 𝛼′ were varied during the study.

On the other hand, the inner diameter 𝐷0, the mating angle 𝛼 and the separation angle 𝛼′ were varied during the study. The material used for this study was polyamide 12 (nylon), so the mechanical properties such as E and μ were also set during the whole study.The only parameter which was assigned arbitrarily was the permissible deflection δ; it has been set to be 1 mm. This parameter indeed depends on the permissible strain in outer fibers ϵ which value was not known. A finite element simulation would have been necessary to obtain the value of ϵ.Once all the parameters had been fixed, fifteen snap-fit pairs including fifteen top parts and bottom parts were designed.

An example of a pair is shown in Figures 4 and 5.Figure 4: Top partFigure 5: Bottom partFigure 6 shows the three geometrical parameters that have been varied for this study.𝐷0𝛼′𝛼Figure 6: Detailed drawing of a pairA 3D Systems HiQ Sinterstation with StablTemp capability was used to fabricate the samples. In this study, six different separating angles have been tested (30°, 40°, 50°, 60°, 70°, 90°), as well as, five different mating angles (30°, 40°, 50°, 60°, 70°) and four different inner diameters (4 mm, 6 mm, 8 mm, 10 mm).The loading tests made on the pairs of samples were tension and compression tests, performed using a 500 N capacity Instron universal tester. The results were recorded and compared to the values obtained with the derived equation.

The results were recorded and compared to the values obtained with the derived equation. Clamps were used during the tests to achieve best possible colinearity between the top part and the bottom part.Figure 7: Tension test of a pairResults and DiscussionsAccording to Figures 8 and 9, the values found with the derived equation seem to be close to the results found experimentally, although the experimental curves do not exactly follow the ones plotted with the analytical results. This can be explained because of the difficulty to get a perfect collinearity between the two parts during the tests.

This can be explained because of the difficulty to get a perfect collinearity between the two parts during the tests. Moreover, obtaining an identical contact at the same moment for the four branches is extremely complicated when the force is applied on the snap-fits which resulted in some non-representative results.800.00700.00600.00500.00D0 = 10 mm α' = 45°)N( e c r o F400.00300.00Fm_calc (N)200.00Fm_exp (N)100.000.003040506070α (°)Figure 8: Mating force as a function of the mating angle100.0090.0080.0070.00D0 = 10 mm α = 45°)N60.00( e c r o F50.0040.0030.00Fd_calc (N)20.00Fd_exp (N)10.000.003040506070α' (°)Figure 9: Dismounting force as a function of the separation angleThe gap present for α = 70° is especially large, because the sample broke during the compression test.

This can perhaps be avoided with a larger contact surface between the two parts.Figure 10 shows a sizeable difference when comparing the analytical and experimental results. This is most likely due to the same issue as mentioned for the two graphs above.1600.001400.00α = α' = 45°1200.00)1000.00N(t r o f f E800.00Fm_calc (N)600.00Fd_calc (N)400.00Fm_exp (N)Fd_exp (N)200.000.0046810D0 (mm)Figure 10: Mating / Dismounting force as a function of the inner diameterOverall, the force values obtained for tension tests and compression tests where α = α' are not the same. Indeed, the force value should be the same whether it is in tension or in compression, according to the derived equation.A possible explanation of this observation is the presence of some degree of buckling during the compression tests when the load is applied on the top part to mate with the bottom part.

This results in the bowing of the top part’s four branches during this phase and thus more load is necessary to counter this effect.However, this model has its limits. In fact, when α or α‘ reach 80°, (value that has been found by interpolation between 70 degrees and 90 degrees), instead of having higher force values, lower values have been obtained with the derived equation. The reason why the force . In this case, both theμ+tan αdecreases comes from this part of the derived equation :1−μ∗tan αnumerator and the denominator tend quickly to infinity which creates an indeterminate form. μ , it follows that the obtained coefficient is inferior to theμ+tan α 1−μ∗tan α ~ −1Thus, limα→90 previous ones starting from α = 80°.ConclusionsAn experimental study has been done on polyamide 12 snap-fits, fabricated with a Selective Laser Sintering machine, by performing compression and tension tests on fifteen paired samples.

The results obtained were compared to analytical results obtained using a derived equation. The geometrical parameters chosen to make the comparison were the mating angle α, the separating angle α’ and the inner diameter 𝐷0. After analyzing the results, work must be into account the buckling effect during the done on the current model to take compression/mating and to allow to understand more fully why the equation seemingly breaks down for higher angle values.Despite this fact, this model seems to be accurate for angles between 30° and 70°. However, perfect collinearity and identical contact for the four top part branches are required during the tests to obtain representative results. All in all, additive manufactured snap-fits can be integrated in systems during CAD to print them directly on the main part.

All in all, additive manufactured snap-fits can be integrated in systems during CAD to print them directly on the main part. This provides a means to attach components to one another, for plastic parts at least.AcknowledgmentsOne of the authors (KT) performed this work while on an undergraduate internship at the University of Texas at Austin.

perimental Study of the Cooling Characteristics of Polymer Filaments in FDM and Impact on the Mesostructures and Properties of Prototypes.Q. Sun, G.M. Rizvi, C.T. Bellehumeur 1* and P.Gu2 1Department of Chemical and Petroleum Engineering 2Department of Mechanical and Manufacturing Engineering University of Calgary, Calgary, Alberta, T2N 1N4 CanadaReviewed, accepted August 19, 2003AbstractThe bonding quality among polymer filaments in the fused deposition modeling (FDM) process determines the integrity and mechanical properties of the resultant prototypes. This research investigates the bond formation among extruded acrylonitrile butadiene styrene (ABS) filaments in the FDM process. Experimental measurements of the temperature profiles were carried out for different specimens and their effects on mesostructures and mechanical properties were observed. Models describing the formation of bonds among polymer filaments during the FDM process are discussed.

Models describing the formation of bonds among polymer filaments during the FDM process are discussed. Predictions of the degree of bonding achieved during the filament deposition process were made based on thermal analysis of extruded polymer filaments. The bond quality was assessed based on the growth of the neck formed between adjacent filaments and their failure under flexural loading. Further experimental work is underway to assess the validity of the proposed models.1. IntroductionRapid prototyping (RP) is one of the fastest growing manufacturing technologies for fabrication of cost effective models, prototypes and one of a kind parts. It is being applied in a very large number of fields, such as aerospace, automotives, architecture, and medicine. RP technologies provide the ability to create almost completely finished products from CAD files.

RP technologies provide the ability to create almost completely finished products from CAD files. Moreover, it has been shown that some RP techniques have the potential to fabricate parts with locally controlled properties (porosity, density and mechanical properties) [1,2]. The FDM machine is basically a computer numerically controlled (CNC) gantry machine, carrying double miniature extruder head nozzles. In the FDM process, parts are fabricated by extruding a semi- molten filament through a heated nozzle in a prescribed pattern onto a platform. As the material is deposited, it cools, solidifies and bonds with the surrounding material. The fabrication of parts with locally controlled porous structure and mechanical properties can be achieved by varying the deposition strategy, deposition orientation and other process conditions.

The fabrication of parts with locally controlled porous structure and mechanical properties can be achieved by varying the deposition strategy, deposition orientation and other process conditions. However, for directly producing functional parts, fundamental understanding of mechanical properties of FDM parts with respect to fabrication process parameters is essential.Corresponding Author: C.T. Bellehumeur, cbellehu@ucalgary.ca, tel.: (403) 220-8804, fax: (403) 282-39453131FDM prototypes are composites of partially bonded filaments and voids. The bonding quality among filaments in FDM parts is an important factor in determining the integrity and mechanical properties of resultant prototypes. The formation of bonds in FDM process is driven by the thermal energy of semi-molten material. The quality of bonds formed between individual filaments depends on the growth of the neck formed between adjacent filaments and on the molecular diffusion and randomization at the interface.

The quality of bonds formed between individual filaments depends on the growth of the neck formed between adjacent filaments and on the molecular diffusion and randomization at the interface. The bond formation process can be viewed as a sintering process for which the wetting phenomenon is of importance. It can also be modeled following approaches similar to those used to describe polymer welding, where the issue of molecular diffusion dominates.Studies in FDM have primarily been directed towards the development of new materials or techniques for material deposition [3-5]. The mechanical properties of filaments and FDM parts were also investigated by a number of researchers [6-8]. However, considerable research efforts are still needed to fully understand the mechanical properties and their relations to the fabrication processes. In particular there is a lack of literature on experimental determination of the thermal profiles of the parts built using FDM.

In particular there is a lack of literature on experimental determination of the thermal profiles of the parts built using FDM. This data is especially important, as it is indicative of the bond formation between adjacent filaments. This paper provides the results of experimental determination of the thermal profiles of some simple shapes produced using the FDM process and their effects on the bond formation in terms of neck growth between adjacent filaments and the intermolecular diffusion at the interface and their evaluation in regards to the mechanical properties of the parts.2. Experimental WorkThe FDM 2000 from Stratasys Inc was used in this study. A commercial acrylonitril butadiene styrene (ABS P400) was used. The material was heated to 270°C and extruded through a nozzle onto a platform. The envelope temperature was set to 70°C. All parts were built using tip 12 with zero air gap.

The material was heated to 270°C and extruded through a nozzle onto a platform. The envelope temperature was set to 70°C. All parts were built using tip 12 with zero air gap. In addition, parts produced in this study consisted of unidirectional specimens to enhance the effects of the parameters being studied. The parts produced for mesostructure characterization and for mechanical testing were built with the normal five layers of support material. As the support material may have different characteristics, the parts generated for the measurements of the filament’s temperature profile were built without the support layers.The temperature profiles of extruded filaments were monitored using 0.003" K type thermocouples. The thermocouples were imbedded in the foam of the base plate of the FDM machine.

The thermocouples were imbedded in the foam of the base plate of the FDM machine. The measurements were recorded and analysed through a high-frequency analog- digital converter (1000 readings per second), amplifier, data acquisition card 6024E and software Labview 6.0 on a PC.The samples collected from the parts were sectioned using a Leica RM2165 microtome. The samples’ cross-sections were viewed under optical microscopy. Pictures of these cross-sections were taken with a CCD camera mounted on an Olympus BX60 microscope and features were analyzed using the image analysis software Image-Pro®.3142Failure LineFigure 1: Flexural strength test sampleThe flexural strength testing of the specimens was carried out on the ATM testing machine (Model TTS Series). The tests conformed to ASTM standard D 1184-98 for flexural strength of specimens were laminated sheets bonded with glue.

The tests conformed to ASTM standard D 1184-98 for flexural strength of specimens were laminated sheets bonded with glue. The constructed so that the failure occurs in the bonds between adjacent filaments as shown in Figure 1. One deviation from the standard was that failure occurred between filament bonds in the same layer instead of the bonds between adjacent layers. The test also differed from the standard in that 12 layers of laminates were used instead of layers, so that parts with reasonable the stipulated eight dimensions could be tested.3. Thermal Processes in FDMThe bonding between individual roads of the same layer and of neighboring layers of FDM parts is driven by the thermal energy of the semi-molten material that is extruded from the FDM head. The temperature history of interfaces plays an important role in determining the bonding quality.

The temperature history of interfaces plays an important role in determining the bonding quality. When the filament is deposited and is in contact with surrounding material, the interface’s temperature is well above the material’s glass transition (Tg). This condition favors molecular diffusion at the interface and the development of adhesive bonds.Several heat transfer models describing the FDM process have been proposed in the literature. Yardimci [9] developed a family of numerical models for the fused deposition ceramic (FDC) processes. A similar analysis method is applicable to the FDM process that uses thermoplastic polymer as a raw material. Rodríguez-Matas [10] presented a two-dimensional heat transfer model in which the filament is laid on a stack of other filaments. The model assumes constant heat convection coefficient and perfect contact among stacked filaments.

The model assumes constant heat convection coefficient and perfect contact among stacked filaments. The results of the analysis show that the interface temperature rises above the glass transition temperature (Tg) immediately upon extrusion of the filament and stays above for a short period of time. In Rodríguez’s analysis, the variations in the temperature across the filament thickness are taken into consideration, which allows for consideration of the boundary conditions at the interface. The model, however, neglects all contact resistances between the filaments. Based on this 2D model, the conduction heat transfer with the foundation is dominant over the convection heat transfer with the environment. Li and co-workers [11-12] used the lumped-capacity analysis for modeling the cooling process of the extruded filament, because the diameter of the extruded filament is fairly small and a uniform temperature distribution throughout the filament cross- section can be assumed.3.1.

Temperature Profile of a Single FilamentExperimental measurements performed on a single filament deposited on the base plate foam platform were used to assess the validity of the two models described above (Figure 2). The material properties were selected based on the work presented by Li et al. [12]. The measured data is in general agreement with these models, lying somewhere in between the two predicted values. It is interesting to note that at the higher temperatures prevailing in the initial3153)C ° (e r u t a r e p m e T280230180130Glass transitionExperimental datah=75W/m2.K 2D Model [11-12]h=75W/m2.K LC Model [10]stage, the measured temperature profile exhibits more similarity to the lumped capacity model predictions, while at lower temperatures the measured data tends to follow the 2D model predictions [10].

The filament reached Tg in about three seconds, which is not very different from the predictions of both models.80Figure 2: Single filament cooling profile where h denotes the convective heat transfer coefficient.02Time (s)463.2. Multi Layered Specimen: Temperature Profiles and Mechanical TestsTwo sets of specimens were constructed to carry out mechanical tests in order to assess the bond strength between the filaments. The dimensions of the test specimens were 3.1 mm x 19.1 mm x 31 mm, and all the filaments were oriented perpendicular to the longitudinal axis of the parts for the tests so that failure would occur in the bonds between the adjacent filaments of the lower layer as shown in Figure 1. In the first set, the specimens were built individually, whereas for the second set, three specimens were cut from one longer piece (Figure 3a).

In the first set, the specimens were built individually, whereas for the second set, three specimens were cut from one longer piece (Figure 3a). The temperature profiles at the base of each representative specimen were also recorded and are shown in Figure 3b.150140Glass transition130Small part (31mmx19mm))C ° (e r u t a r e p m e T120110100Longer part (100mmx19mm)9080700500100015002000Time (s)Figure 3a: Different specimens for flexural strength testsFigure 3b: Comparison of the temperature profiles for different specimens3164One must note that the highest temperatures recorded in these profiles were lower than that measured for a single filament. This was due to the fact that a larger set of data (1000) was averaged to get one record per second, compared to 100 records per second for the previous case. Therefore, the higher values were averaged with lower values to give smaller peaks.

Therefore, the higher values were averaged with lower values to give smaller peaks. Secondly, the models discussed in the previous section are only valid for single roads and cannot be applied to the situation represented in this case, where conduction is the dominant mode of heat transfer from the adjacent and upper layer filaments. It can be seen that for the specimens built individually (first set) the temperature at the base of the part remains above the glass transition temperature (Tg) after the 4th layer is built and continues to be high until the parts are completed. For the second set of specimens, the temperature reaches a level above Tg for only a short duration during the deposition of each layer of filaments. Therefore, it is to be expected that the individually made specimens will be stronger than the specimens cut from longer pieces. This expectation is indeed borne out by the mechanical testing and the results are shown in Table 1.

This expectation is indeed borne out by the mechanical testing and the results are shown in Table 1. The flexural strength of the individually made specimens was 2810 kPa, whereas, it was 2277 for the specimens cut from the larger part. The higher temperature profile, which gave rise to higher flexural strength, was the result of the shorter traverse of the FDM head for making one layer of the individually built specimen compared to the three times larger traverse to build the longer part from which three specimens were cut. This resulted in better bonding between the adjacent filaments and consequently the individually built parts exhibited higher failure loads.The test data exhibited a much larger standard deviation in the results of the specimens cut from the longer parts (Table 1). This deviation could be the result of two possible factors. The first was that a number of larger parts exhibited miniature fault lines between adjacent filaments.

This deviation could be the result of two possible factors. The first was that a number of larger parts exhibited miniature fault lines between adjacent filaments. One possible cause of these faults could be the wear in the mechanical system of the FDM machine. The obviously faulty samples were discarded on visual inspection, but if the fault was in the inner layers, or too small, it would escape visual inspection. Another factor was that the temperature profile was observed to change with variations in the location of the parts built. This would cause variation in exposure temperature/time of the parts being built.

This would cause variation in exposure temperature/time of the parts being built. Further experimentation is required before drawing any definite conclusions.Part typeFlexural strength (kPa) Standard deviation (kPa) Total samples usedCut from longer piece2277.0105.98Individually built2810.087.24Layer undergoing failure Flexural strength (kPa) Standard deviation (kPa) Total samples usedBottom2277.2105.98Top2153.665.95Table 1: Comparison of the flexural strength for individual parts of different sizeTable 2: Comparison of the flexural strength for failure in the top vs. bottom layers of the specimens cut from the longer partThe temperature profile of the longer part shown in Figure 3b also suggests that the top layers of the specimen are exposed to shorter durations of temperature above Tg compared to the lower layers. This would indicate that the lower layers of the specimen should have higher bond strengths than the upper layers.

This would indicate that the lower layers of the specimen should have higher bond strengths than the upper layers. To verify this, a number of samples were prepared and tested so3175that the failure would occur either in the top or in the bottom layer of the specimens, and the results are shown in Table 2. Indeed, the bottom layers did exhibit higher bonding strength between the adjacent filaments. The flexural strength of the specimen when the fracture was caused to occur in the bottom layer was 2277 kPa against a value of 2153 kPa when the fracture was caused to occur in the top layer. Again the standard deviation was quite large, since the specimens were cut from the longer parts as already discussed above.3.3. Effects of Prolonged Exposure to Higher Temperatures on the MesostructureIt was generally observed that the temperature profile of the bottom layers of the parts remains at temperatures higher than Tg for a longer period of time compared to the upper layers.

Thus, different neck growth between adjacent filaments can be expected in the top and bottom regions. However, this phenomenon was not readily observable in a twelve-layer specimen. Consequently, thirty-layer specimens were prepared. Figure 4 shows a magnified picture and the the temperature profile for these specimens. As can be seen, after the first few layers, temperature was above Tg until the completion of the part. The voids in the lower region are clearly smaller than in the upper region indicating larger neck growth in the bottom of the part. The dimensionless neck growth in the top few layers is 0.187 compared to 0.310 in the bottom layers. One possible explanation for the larger neck growth could be the longer diffusion time as discussed in the later sections dealing with modeling.

One possible explanation for the larger neck growth could be the longer diffusion time as discussed in the later sections dealing with modeling. Other explanations could be the effect of gravity or the effect of continued downward pressure from the FDM head while the part is being built.Top layers230210Bottom layers)C ° (e r u t a r e p m e T190170150130110Glass transition30 layers90 70010002000 Time (s)3000Figure 4: Mesostructure and temperature profile of a 1.5” x 1.5” x 30-layer part3.4. Effects of Filament’s Deposition Strategy on Temperature Profiles and MesostructuresThe impact of the deposition strategy on the parts’ characteristics was examined in this work. Figure 5 shows the temperature profile at the base of a part with dimensions 200 mm x 12 mm x 2.8 mm in which the filaments were deposited in the longitudinal direction.

Figure 5 shows the temperature profile at the base of a part with dimensions 200 mm x 12 mm x 2.8 mm in which the filaments were deposited in the longitudinal direction. After the second layer is formed, the temperature at the base remains above Tg for the duration of the build time. Figure 6 shows the temperature profile at the base of a part with dimensions 130 mm x 28 mm x 2.8 mm in which the filaments were deposited in the lateral direction. Due to the short path3186length, the temperature reaches much higher values as the FDM head approaches the thermocouple, but later drops below the Tg limit as the head recedes. In this case the part is exposed to higher temperature for a longer time. Table 3 summarizes the results obtained for the dimensionless neck formed between adjacent filaments.

In this case the part is exposed to higher temperature for a longer time. Table 3 summarizes the results obtained for the dimensionless neck formed between adjacent filaments. The lateral specimen exhibits a higher dimensionless neck growth, and consequently has a lower void fraction due to longer exposure to higher temperatures.)C ° (e r u t a r e p m e T170 150 130 110 90 70Glass transitionExperimental data)C ° (e r u t a r e p m e T170 150 130 110 90 70Glass transition Experimental data050010001500050010001500Time (s)Time (s)Figure 5: Temperature profile for a longitudinally built partFigure 6: Temperature profile for a laterally built partDimensionless Neck growth Void fraction (%)Longitudinal parts 0.1910.091Lateral parts 0.2330.081Top layers (30-layer part) 0.187--Bottom layers (30-layer part) 0.310--Sintering model prediction0.152--Table 3 Dimensionless neck growth and void fraction for various parts4.

Prediction of Bond FormationThe formation of bonds between polymer filaments in the FDM process can be described as shown in Figure 7. The cross-sections of filaments are idealized as circles in the figure. The first step of the process is the establishment of interfacial molecular contact by wetting. The molecules then undergo motions toward preferred configurations to achieve the adsorptive equilibrium [9]. Molecules diffuse across the interface, forming an interfacial zone, and/or react to form primary chemical bonds across the interface. The randomization can be reached only after extensive inter-diffusion of chain segments under critical conditions. These processes can be interpreted either in terms of sintering of adjacent particles or inter-diffusion of molecules at the contact interface, both of which occur at elevated temperatures.

These processes can be interpreted either in terms of sintering of adjacent particles or inter-diffusion of molecules at the contact interface, both of which occur at elevated temperatures. Predictive models, for both sintering and diffusion, which utilize the measured temperature profiles, are discussed in the following sections.7319(cid:1)(cid:1)Figure 7: Schematic of bond formation between two filaments4.1. Sintering ModelAnalytical models describing the sintering rate that occur, due to Newtonian and viscoelastic flows, in various polymers have been proposed and assessed by Hornsby and Maxwell [13], Mazur and Plazek [14], and Bellehumeur et al. [15-16]. In this work, the model proposed by Pokluda and co-workers [17] has been used.θd dt=Γoaµ2−53 ( 1cosθ−cossin )( 1θ( −θ 2+cosθ cos ) 3 1θ)13(1)−ay1=θwith.

The model is valid for the coalescence of two particles and predicts the neck growth between the particles as a non-linear function of time t, material viscosity µ, initial particle radius ao, and surface tension Γ. Details on the mathematical development and assessment of the model have been discussed elsewhere [11-12].sinis used to determine the neck growth formation between the sintering particles, thus providing information about the degree of wetting achieved at the filaments’ interface, which is an indication of the neck growth. Previous work [1] has shown that the neck growth becomes negligible at temperatures below 200°C for ABS P400. This temperature is to be referred to as the critical sintering temperature for ABS P400. Quantitative predictions about the degree of bonding achieved during the filament deposition process were made based on the lumped-capacity heat transfer and Newtonian sintering models.

Quantitative predictions about the degree of bonding achieved during the filament deposition process were made based on the lumped-capacity heat transfer and Newtonian sintering models. The selection of model parameters was based on criteria presented in reference [12]. These results were found to be only in general agreement with the experimental observations (Table 3). Quantitatively, the difference might due to simplifying the cross-section shape of the filament (circle versus ellipse), neglecting gravity force in the sintering model and neglecting elastic deformation that may occur at temperatures below the critical sintering temperature.The sintering model4.2 Diffusion ModelThe theory of intermolecular diffusion across the interface under isothermal conditions is well established in literature. The isothermal models work quite well for quasi-isothermal processes such as autoclaving.

The isothermal models work quite well for quasi-isothermal processes such as autoclaving. However, for on-line consolidation type processes such as FDM, where thermal transients are present at multiple time scales, a non-isothermal model is needed. A3208model put forwarded by Fang [18] for the healing process under non-isothermal conditions was used to predict bond strength. The model considers a temperature dependent welding time and describes the bond strength as a function of the temperature history. The non-isothermal degree of healing evolution with time is developed starting from a fundamental formulation of the reptation of polymer chains and is given by the following expression:tD )( h=σ σ ∞=(cid:6) (cid:4) (cid:5)t(cid:7)01 Tt )( wdt(cid:3) (cid:1) (cid:2)4/1(2)where tw(T) is the temperature dependent welding time, (cid:0) is the strength of the fully healed interface, and (cid:0) is the stress for the FDM part.

For a given temperature history, and a given temperature dependence of the welding time, the degree of healing can be evaluated using equation 2. The respective temperature profiles in conjunction with the extruded ABS P400 welding time, as determined by Rodríguez [8], were used to quantitatively predict the bond strength of the parts produced for mechanical testing. The welding time used is only valid in a very narrow temperature range and needs to be re-examined at higher temperatures in future work.(cid:2)=/ where Qd = 388.7 kJ/mol and C = 1.080×10-47 s.tCexp(Q dwRT)(3)The model described by equation 2, was implemented numerically to predict the degree of healing as a function of time. The numerical integration was carried out until the degree of healing reached a value of unity, or until the specified time was completed, whichever occurred first.

The numerical integration was carried out until the degree of healing reached a value of unity, or until the specified time was completed, whichever occurred first. Furthermore, for amorphous materials the healing mechanism is active only as long as the material time when the temperature was above the glass transition point was taken into consideration.temperature is above its glass transition point, hence only that0.5The modeling results are shown in Figure 8. From Figure 3b it can be seen that the individually produced parts stay above glass transition temperature longer and consequently have a higher degree of healing with the resultant higher bond strength. Figure 8 shows a drop of 28% in the degree of healing for the longer part from which the three specimens were cut. However, the drop in the load bearing capacity of these specimens was only about 10%.

However, the drop in the load bearing capacity of these specimens was only about 10%. The difference can be attributed to the available welding time data, which covers only a narrow temperature range, but is extrapolated for a wider temperature range, causing errors. As an example, the extrapolated welding time for a temperature of 170°C was 0.01 seconds, which is obviously not correct. For the shorter part, the time spent at higher temperatures was greater, resulting in over estimation of the degree of healing.g n i l a e h0.40.3f oSmall parts0.2e e r g e D0.1Longer parts0300 Time (s)5001000400200Figure 8: Degree of healing of FDM parts with different temperature histories32194.3 Model ComparisonUnder the conditions selected for the processing of ABS P400, sintering has a significant effect on bond formation, but only for the short duration of time that the temperature is higher than 200° C.

Experimental observations showed that the filament temperature is below this temperature and above the glass transition point for a considerably longer period of time where the bonding is driven by the molecular diffusion at the interface. Therefore, the sintering model can be applied to the initial period when the filament temperature is above 200° C and the diffusion model is more appropriate for predicting subsequent bond strength development.5. ConclusionsIn this paper, on-line measurements of the cooling temperature profiles of extruded filaments in the FDM process were carried out. It was found that most theoretical models for the cooling profile of the extruded filaments fail to predict the temperature history of filaments embedded in a multi-layered part. Experimental measurements of the temperature profiles for different specimens and their effects on the mesostructures and mechanical properties were discussed.

Experimental measurements of the temperature profiles for different specimens and their effects on the mesostructures and mechanical properties were discussed. Based on the heat transfer analysis and the sintering experimental data, the prediction of the growth of the neck formed between adjacent filaments was evaluated using the sintering model. A non-isothermal diffusion model was used for the prediction of the bonding evolution at the interface between filaments due to molecular diffusion.The future research will be focusing on obtaining accurate estimates for the non- isothermal diffusion model and generating predictions of bond strength development between filaments in FDM parts. Parametric studies will be conducted to investigate the effects of various process conditions on temperature histories and on the evolution of mechanical properties with time.6.

..4""""' ....... _1!1.1!1......... ~"'1!1.... ,.,............ ,iI' ... ,. .....,"',. of the _n1r'01ll"ll"lllO'lrDll11"'~",uu.",_ UsingEdward Gargiulo**,MichaelINTRODUCTIONphotopolymer solid imaging process for """"'''''.JUA.UAjiO, a laser beam to sequentially draw or ...., liquid polymer.[l] With an "".."""'v. lU ""United States alone, there is considerable n .. ,....,n""'''U'''AA'"''''''jiO,.I.''Ap., goals of rapid prototyping is the \.!LAIl-AAAAU.""", dimensions of a part built on abalanced to producedto many problemsvertical wall post-cure distortion"''''''ll''"I.,.t"" polymerization is shrinkageattendantstresses often result in distortion of the '.II'rlV._'n' ....'....A..,,'~'VA "",AU' " A has been discussed by several authorsmay be defined as any stresses residual to.."V "A" extensively and has """1"'''"1'''''''U ..t2JI.l\JI.ll..." and cantilever length..1"I'"IlJ...

, ...., ...U'IU''U''''''''''''''' and various parameters AU....'AAIl-LAL...".U,curl distortion still needs towas to experimentallywarp deformation. A U....,"" ...jiO,All...., ......u ...~JV ... I,...... " ....V of the parameters.178v1·.11...THEORYcourse of preparing a part for stereolithography as many as twenty parameters mayselected that will govern the imaging process and the qualities of the final object.VOt.:l........... 'W-lIo&JL parameters were selected for inclusion in this study:Layer Thickness, Hatch Style, Hatch Cure Depth, Hatch Spacing, and Fill Cure Depth.Three-dimensional objects are "sliced" into series of constant thickness, two-dimensional to be printed in solid imaging processes. Layer thicknesses are chosen to balance the need to ,"",11'\11,"",,1'713 geometric errors in the build direction caused by the inevitable "contouring" of curved surfaces against the desire to achieve a reasonable build time.

A few practitioners regularly build parts with thicknesses as thin as 0.0025 inches, but most select a thickness near 0.010 inches.Each two-dimensional layer is imaged on the photopolymer surface with a specific hatch by the, user. Hatching is used to photopolymerize the liquid in the interior regions of style a part. hatch style affects the sequence and amount of polymer solidification in the part and therefore its physical properties and internal stress distribution. Several hatch styles have been develoPed for photopolymer solid imaging, but only two will be considered here.TRIHATCH style images vectors in three directions (0, 60, and 120 degrees relative to the x-axis) to form equilateral triangles, Figure 1. The vectors are imaged without offsets between layers and with cure depths greater than the layer thickness to assure layer to layer adhesion. Up and down-facing skin fills trap an amount of liquid in the triangular prisms that are formed.

Up and down-facing skin fills trap an amount of liquid in the triangular prisms that are formed. The volume of residual liquid resin is related to the spacing and linewidth of the hatch vectors. TRIHATCH imaging was the first commercialized with acrylate resins and systems withlow power and short life He..Cd lasers to provide a fast and efficient imaging process. TRIHATCH imaged parts exhibit low curl, but may suffer from POstcure warping and swell distortion.Weave hatch styles image parallel vectors first in the x-direction, followed by parallel vectors in the y-direction. Hatch spacing is selected to be greater than the cured linewidth so the parallel vectors do not interact. Cure depths may be chosen to be less than or greater than the layer thickness depending on the hatch spacing.

Cure depths may be chosen to be less than or greater than the layer thickness depending on the hatch spacing. The STARWEAVE variation of this style,Figure 2, vectors in sequential layers (STagger), alternates the order of x and y vector writing offsets (Alternating), and alternately stops the vector imaging shoftofthe opposing borders (Retraction). The cure depth in this hatch styleis usually less than the layerthickness. The principal advantages Weave imaging styles are the elimination of swell andpost..cure warp distortions. Curldistortions are not universally improved with these stylesdistortion can occur in aU rapid-prototyping methods that build parts in successive the solidifying material undergoes shrinkage. This shrinkage causes distortion and .........,..........'...... stresses. [5] Figure 3 shows the sequence of steps leading to curl distortion. When a single layer is ,first imagedonthe,liquid.surface,itisfree to shrink without inducing stresses.

stresses. [5] Figure 3 shows the sequence of steps leading to curl distortion. When a single layer is ,first imagedonthe,liquid.surface,itisfree to shrink without inducing stresses. However, the second and subsequent layers that are drawn are each bonded to thelayer below that has already shrunk. If there is shrinkage of these upper layers after they have become bonded to below, a bending moment is introduced that can cause upward displacement of the unsupported ends of the layer.

If there is shrinkage of these upper layers after they have become bonded to below, a bending moment is introduced that can cause upward displacement of the unsupported ends of the layer. Curl distortion is typically measured in terms of a curl factor [4] defmed as the vertical distortion distance, h, divided by the length of the free layerFigure 1: TRIHATCH Imaging StyleE!JGm [(JWSTagJEre::l Wf!fiNeAlta"nateSEqUenCingReta::'ld hathFigure 2: STARWEAVE Imaging StyleIrraJe LayerLayEr StrinkcgeIrm~ LayErL"'I-l I'" ~~ hLayEr StrinkcgeFigure 3: Layer by Layer Build ProcessEXPERIMENTThe purpose of our experiment was to statistically determine the relative importance of various build parameters on the curl distortion. The parameters studied were hatch type, layer thickness, hatch spacing, hatch overcure and fill cure depth. The experiment was divided into four sets of parts.

The parameters studied were hatch type, layer thickness, hatch spacing, hatch overcure and fill cure depth. The experiment was divided into four sets of parts. Each set consisted of one hatch type and one layer thickness with three different179avalues for each of the three remaining parameters. Thus there were 27 parts for each of the four sets for a total of 108 parts built. Table I summarizes the experiments.Table I: Ranges of Experimental ParametersExperiment Hatch TypeNumberHatch Thickness SpacingLayermilmilHatch Overcure milFill Cured Depth mil1 2 3 4TRIHATCH TRIHATCH STARWEAVE STARWEAVE10 5 10 510,30,50 10,30,50 10,15,20 10,15,202,2,6 -2,2,6 -2,-1,0 -2,-1,05,15, 25 5,15, 25 0,12.5,25 0,12.5,25(1 mil = 0.001 inch)Figure 4 shows the solid model geometry used to create the test pieces. Since the geometry is identical for all the pieces, the curl distortion is proportional to the chord height, h.

Since the geometry is identical for all the pieces, the curl distortion is proportional to the chord height, h. A dial indicator depth gage was used to measure that height for each experimental piece and that value was recorded as a measure of the curl distortion.Figure 4: Test PartThe two hatch types chosen were TRIHATCHand STARWEAVETM. The TRIHATCH pattern leaves a large portion of the part in the liquid state. Post curing must be used to solidify the honeycomb trapped volumes and results in internal stresses. The STARWEAVE hatch style was d~veloped to reduce those internal stresses. The pattern is formed by creating orthogonally alternating layers of fingers that are not all attached to their respective ends. Furthermore, the hatch overcure is selected so the layers are not completely attached.

Furthermore, the hatch overcure is selected so the layers are not completely attached. Since curl distortion is related to the extent of shrinkage after contact with the previous layer, the delayed shrinkage is expected to reduce curl distortion.Layer thicknesses of 5 and 10 mils were chosen for this study. These values are commonlyused by many users and they are usually selected to meet l.iCcuracy and build speed requirements.Hatch spacing is the distance between parallel vectors used to hatch the interior of the part. If the hatch spacing is very small, the solidifying vectors will overlap causing a completely solid layer. Large hatch spacings anow liquid polymer to be trapped in the part to be solidified in the180postcureing operation. Values were selected for each hatch style to avoid complete vector overlap and show the postcure effects.Hatch overcure is the depth that one cured vector "string" pierces into the lower adjacent layer.

This is what keeps the individual layers connected together to form aco.mplete. part. Although the major portion of the STARWEAVE vectors are not intended to overlap,multiple exposures at vector intersections are expected to "nail" the layers together at/iPQints. The TRIHATCHconstruction superimposes vector scans creating a line connection betwe~nlayers. Hatch overcures were selected to try to make sure the parts had sufficient layer over prevent delamination and give good "green" strength. Nominal hatch overcures of 6 mils are usually recommended forTRIHATCH and a mil overcures (1 mil undercure) is a common starting point for STARWEAVEeFill cure depth is tbe depth of the solid layers formed on the upper and lower faces of the fill s aresolid. This holds the remaining liquid inside the part for subsequent post-curing. required for the TRlHATCH style, but are sometimes eliminated in STARWEAVEeMany other parameters need to be defined for a complete build.

required for the TRlHATCH style, but are sometimes eliminated in STARWEAVEeMany other parameters need to be defined for a complete build. Thesewere all held constant The Appendix gives the complete setup for the 3D ..... 'c·r""'n.. " to create the parts.RESULTSThe following tables show the chord height for each of the pieces made as measured by a dial depth gage. Positive values indicate an upward curl as shown in Figure 3, while negative values indicate sag. An "X" indicates that the style was not able to produce a viable solid (e.g..U .. ":;Olb............,..v,u""strength or delamination). All dimensions are in mils (0.001 inches).Tablesthrough V show a large number of failures and thus theremake a complete statistical evaluation. However, some qualitative observations. can be the TRIHATCH with 5 mil layer thickness, one can estimate that the 30 and 50 mil hatch SpaClnj~S were tooapart, with 30 mils close to the upPer limit.

For the STARWEAVE style, the curl values are all fairly large. It seems that the range of parameter values chosen were not appropriate for the STARWEAVE pattern. More experiments need to be performed with different parameter values or perhaps different parameters. This indicates more understanding and practical experience with the TRIHATCH pattern.The only set to give enough data was the TRIHATCH with 10 mil layer thickness. Qualitatively, Table IT indicates curl minimization near the 30 mil hatch spacing which matches our experience. To study the data, we first performed a three-way layout with one observation per cell analysisvariance.

To study the data, we first performed a three-way layout with one observation per cell analysisvariance. The three factors were hatch spacing, hatch overcure and fill cure depth:h(i,j,k) = K + A(i) + BG) + C(k)where K is a translational constant and A, B, C represent the effects of the three parameters: hatch spacing (i=I,2,3 corresponds to 10,30,50 mils), hatch overcure G=1,2,3 con'esponds to -2,2,6 mils) and fill cure depth (k=I,2,3 corresponds to 5,15,25 mils) respectively. Unfortunately, a simple linear model with just those three variables was able to account for less than 40% of the total variance between the model and the data; R = 0.35, where R is the typical correlation coefficient used in least-squares analyses.

Next, the three-way layout was performed with the model:h(i,j,k) = K + A(i) + BO) + C(k) + AB(i,j) + AC(i,k) + BCG,k)where again, K is a constant and ABC represent the effects of the three parameters, but now AB, AC and BC represent the effects of the combined parameters of hatch spacing with hatch overcure, model hatch spacing with fill cure depth and hatch overcure with fill cure depth respectively. was able to account for over 80% of the variance (R ::::: 0.90) and thus gives a reasonable model estimate of the data. However, an estimate of the standard deviation of the experimental curl observati()ns is fairly large at 5.(j37 mils and indicates the nee~ for more experiments. Nevertheless, we have enough data to determine the relative importance of the various effects (A, C, AB, AC and BC) by testing the hypothesis that all values of any specific parameter are 0; that is the parameter has no effect on the curl.

Table VI shows the probability that the particular parameter has no effect on the curl.182Table VI: Relative Importance of ParametersProbability of NO effectHatch Spacing Hatch Overcure Fill Cure Depth Hatch Spacing x Hatch Overcure Hatch Spacing x Fill Cure Depth Hatch Overcure x Fill Cure Depth2.15% 89.28% 39.02% 28.75% 19.46% 25.69%From this, we can conclude that the most statistically significant parameter is the hatch spacing.Given the results from the analysis of variance, we attempted to fit the expedmental data to aquadratic function of the three variables; that is:h == KO + Kl *x + K2*y + K3*z + K4*x*y + K5*x*z + K6*y*z+ K7*x*x + K8*y*y + K9*z*zwhere the KO, Kl, '" K9 constants were determined by a least-squares minimization of error and the variables x,y,z represent hatch spacing, hatch overcure and fill cure depth respectively.

Unfortunately, this model can only account for a little over 40% of the variance (R == .64) and thus indicates the need for more experiments to get better statistical curl information for the parameters and/or the need for a different mathematical model to predict curl.An additional source of error in these experiments could have been the cleaning of the test pieces. No solvents were used in cleaning the parts because of their known influence on curl and swell. Some parts were very delicate and could have been damaged when their supports were removed.Recommendations for further experiments are to:1. Rebuild pieces to obtain statistical measures of curl for a single set of build parameters,2. Build the TRIHATCH pieces with more hatch spacing values between 10 and 20 mils, and3.

Rebuild pieces to obtain statistical measures of curl for a single set of build parameters,2. Build the TRIHATCH pieces with more hatch spacing values between 10 and 20 mils, and3. Build more STARWEAVE pieces to determine more appropriate build parameters.CONCLUSIONIn this paper we experimentally studied the effect of various parameters on the curl distortion in parts built using stereolithography. From our experiments, only the TRIHATCH build with 10 mils layer thickness had enough data for statistical evaluation. However from observation, our experiments indicate that it is difficult to produce good parts with no fill cure depth More parts need to be built with the STARWEAVE style to determine how those build parameters affect curl. Our results for the 10 mil TRIHATCH show that there are important cross relationships between the hatch spacing, the hatch overcure and the fill cure depth In addition, the hatch spacing is the single most important parameter affecting the curl.

We were not able to generate an acceptable quadratic model for our expedmental data and suggest further experiments.]83REFERENCES1. SLA-250 User Guide, 3D Systems, Inc., April 1991.2. Richter, I and P.P. Jacobs, "The Present State of Accuracy in Stereolithography," Proceedings of IDe Second International Conference on Rapid Prototyping. Dayton, Ohio, 1991.3. Murphy, E.I, Ansel, R.E. and II Krajewski, "Reduced Distortion in Optical Freeform Fabrication with UV Lasers," Stereolithography Users Group Meeting, 1988.4. Jacobs, P.F., RaDid ProtQtyping and Manufacturing - Fundamentals of Stereolithography. Society of Manufacturing Engineers, 1992.5. Weissman, P.T, Bolan, B.A., and Chartoff, R.P., "Measurements of Linear Shrinkage and the Residual Stresses Developed During Laser Polymerization," PrQceedings Qf the Third IntematiQnal CQnference on Rapid ProtQtyping.

 Experimental Study.of the.Relationship between Microstructure and Mechanical Properties of a Ceramic Composite Fabricated by Selective Laser Sintering Uday Lakshminarayan*and H.L.Marcus Center for Materials Science and EngineeringUniversity of Texas, Austin, TX 78712.ABSIRACTAlumina-ammonium phosphate powder blends were processed with Selective Laser Sintering. Ammonium phosphate with a melting point of 190°C, acts as abinder when processed with a laser and holds the alumina (m.p. 2300°C) particles together to form a "green" body. When the green body is heat treated at 850°C for 6hours, ammonium phosphate decomposes evolving ammonia and water vapor. Residual P205 reacts with alumina to form aluminum phosphate.

When the green body is heat treated at 850°C for 6hours, ammonium phosphate decomposes evolving ammonia and water vapor. Residual P205 reacts with alumina to form aluminum phosphate. This results in a composite ofunreacted alumina with a coating of aluminum phosphate around the alumina particles.The variation of compressive strength of these low density ceramic composites wasinvestigated in terms of the particle size distribution the amount of binder in the initial blend. It was observed that the strength depends on the relative density and initial blendcomposition and the critical flaw size. A constitutive equation was formulated to characterize the influence of the relative density, binder composition and the critical flawsize on the strength of the composite.INIRODIJCIIONThe concept of Solid Freeform Fabrication (SFF) to generate parts directly from the CAD data without part-specific tooling has been described earlier [1].

Selective Laser Sintering (SLS), one of the Solid Freeform Fabrication (SFF) techniques being investigated at the University of Texas was first described by Deckard and Beaman [2].The SLS process employs a focused laser beam scanned on the surface of the powder bedto induce selective particle - particle bonding in a thin layer of powder. The part isgenerated by sintering appropriate cross-sections on layers of powder stacked on top ofeach other.

The part isgenerated by sintering appropriate cross-sections on layers of powder stacked on top ofeach other. The unsintered powder in each layer remains in the powder bed duringprocessing to support overhangs and other structures in subsequent layers.Now with DTM Corp., Austin, Texas.44The absence of part-specific tooling in SLS makes the process economicalin small-scale production runs where the fabrication of tools for conventional processesconstitute a substantial fraction of the cost of the parts.Initial applications for SLSinclude design verification models made with engineering plastics such as polycarbonate or nylon and wax patterns for investment casting. In these applications, surface finish and dimensional accuracy are of primary importance and strength is of secondaryimportance.

In these applications, surface finish and dimensional accuracy are of primary importance and strength is of secondaryimportance. However, the application of SLS to generate ceramic and metal parts for functional applications requires a detailed understanding of the factors affecting the strength of the part. In this paper, the effect of initial particle size, the composition of the precursor powder and the intrinsic defects on the fracture behavior of the composite isconsidered.POWDER PRQCESSING BY SLSThe processing of alumina - ammonium phosphate powder blends by SLS was described in detail in references [3] and [4]. During laser processing, ammonium phosphate with a melting point of 190°C melts and holds the alumina particles (m.p. 2300°C) in the desired shape. A Nd:YAG laser (wavelength = 1.06Jlm, Q-switched at 40KHz) focused to a beam diameter of O.5mm was used as the energy source to meltammonium phosphate.

2300°C) in the desired shape. A Nd:YAG laser (wavelength = 1.06Jlm, Q-switched at 40KHz) focused to a beam diameter of O.5mm was used as the energy source to meltammonium phosphate. Laser energy was maintained at 20W and the scan speed was 4cmls. The distance between scans was O.5mm. The layer thickness was 125Jlm. Thebed biasing temperature was room temperature.The "green" part from the SLS processing equipment was fired at 850°Cfor 6 hours. During the secondary heat treatment the temperature was ramped up at 2°C/min with an intermediate soak at 450°C to drive off reaction products. After firing,the parts were cooled in the furnace [3]. During heat treatment, ammonium phosphatedecomposed evolving ammonia and water vapor as by-products. Residual phosphoruspentoxide reacted with the alumina to form aluminum phosphate.

During heat treatment, ammonium phosphatedecomposed evolving ammonia and water vapor as by-products. Residual phosphoruspentoxide reacted with the alumina to form aluminum phosphate. Thus a composite of alumina and aluminum phosphate capable of withstanding intermediate temperatures was obtained in a net-shape by SLS using a blend of alumina and ammonium phosphate as the precursor powders.Compressive strength of the samples was measured on specimens having nominal measurements of O.5"xi "xl It. The specimens were oriented during SLS processing such that the build plane was parallel to the I "xI" cross-section.

The specimens were oriented during SLS processing such that the build plane was parallel to the I "xI" cross-section. The sampleswere ground flat and subjected to compressive stresses in an Instron Universal TestingInstrument with a cross-head speed of 1 mm/s..The loads were applied perpendicular tothe build plane.45INITIAL POWDER ANALYSISX-ray powder diffraction analysis of the as-received precursor powders indicated that the alumina was in its a-form and the ammonium phosphate was in itstetragonal form. Preliminary investigation indicated that a 95.0% pure alumina absorbed laser radiation at 1.06Jlm better than the purer form (99.9%). Hence, the research workreported in this paper was confined to the less pure form of alumina. The major impurities in 95 % pure alumina are 1:Ti02-2.91 %, Si02-0.71%, Fe203-0.35%, MgO-0.14%.Three different alumina particle sizes were tested during our experiments.

The major impurities in 95 % pure alumina are 1:Ti02-2.91 %, Si02-0.71%, Fe203-0.35%, MgO-0.14%.Three different alumina particle sizes were tested during our experiments. The mean and standard deviation of the particle size distribution of the coarse, medium and fine alumina are shown in Table. 1. The ammonium phosphate particle size was maintained below 44Jlm throughout our experiments.Manufacturer's DesignationAverage Size (Jlm)Standard Deviation (Jlm)Duralum 15080.011.8Duralum22069.09.2Duralum4008.75.0Table 1. Mean particle size and standard deviation of alumina particles used inexperiments.RESULTS AND DISCUSSIONThe compressive strength of a porous ceramic composite depends on threemicrostructural factors: 1. porosity (or relative density), 2. binder content and 3. cracksize. Each of these factors is considered briefly here.In general, it has been observed that the strength of a porous material decreases as its porosity increases [5,6,7].

cracksize. Each of these factors is considered briefly here.In general, it has been observed that the strength of a porous material decreases as its porosity increases [5,6,7]. As porosity increases the load bearing area decreases resulting in a reduction of the strength of the material. Balshin [5] andKnudsen [6] have proposed equations 1 and 2 respectively to relate the strength of thecomposite to its porosity.O'=O'o(l-p)m ---- (1) ---- (2) O'=O'oe(-bP)where 0' = Strength of porous solid0'0 =Strength of a fully dense solid (P = 0.0)P = Porosity of solid1Analysis supplied by Washington Electro-Minerals Niagara, NY.46b and m are constants.Both equations are based on the reduction of strength due to a reduction in load bearingarea.

Rice [7] on the other hand proposed an equation of the forma=(1t/2)(11/tatnl)1/2(Ey/(R+L))1/2---- (3)where l1=1tR/(R+A)a=Strength of porous materialR=Pore radiusE=Young's modulus of material y=Surface fracture energyL=Fraction of the grain size A=Pore spacing=4R(1-P)/3P for evenly distributed spherical poresto relate the effect of porosity on the strength of the material in terms of the stressconcentration induced at the tip of the pores. From equations 1-3 it can be generalizedthat a decrease in the load bearing area and stress concentration at pore tip reduce thestrength of the material.Eqn. 4 relates the effect of the binder content to the strength of thecomposite [8]. The G-O.5 term in Eqn. 4 is derived from the Griffith crack criterion forthe fracture of brittle materials. Here it is assumed that the crack size controlling the strength of the material is equal to the grain size.

5).a=Y(Ey/a)0.5---- (5)where a=strength of porous solidY=Geometrical factor E=Young's modulus y=Surface fracture energy4]a=crack sizeAs shown in Eqn. 5, strength varies in an inverse square relationship with the crack length. As a first approximation, it is reasonable to assume that the crack size equals grain size [8]. Based on this assumption, Knudsen [6] formulated Eqn. 7 combining the effects of the pore size and the porosity on strength in one equation.a=kG-ae-bP ---- (6) where a=strength of porous solidG=Grain sizeP=Porosity of solidk, a and b are constants.Based on the discussion thus far, it is inferred that the strength of the alumina - aluminum· phosphate composite depends on its relative density, the binder content and the flaw size. The increase in density with alumina particle size and the amount of binder in the initial powder blend was discussed in an earlier report [1]. Fig.

The increase in density with alumina particle size and the amount of binder in the initial powder blend was discussed in an earlier report [1]. Fig. 1 shows the variation of the compressive strength of the alumina - aluminum phosphate composite fabricated by SLS with the binder content and the alumina particle size. It isseen that at a composition (i.e. 35%), the strength of the material increases with thealumina particle size. This is attributable to two reasons. As the alumina particle sizedecreases, the relative density of the material decreases [9]. In addition, as will be discussed later, very large cracks were observed in compacts with fine alumina (9Jlm).The presence of large cracks lowers the strength of the material.In Fig. 2, the strength of composite is plotted against its relative density showing its linear dependence. Considering composites at a particular relative density,specimens fabricated with the finest alumina had the highest strength.

Considering composites at a particular relative density,specimens fabricated with the finest alumina had the highest strength. It was pointed out in an earlier report [4] that the composites using fine alumina required a higher bindercontent than when using coarse alumina to get the same relative density. The presence ofa large binder content results in a stronger material.

1. Compressive strength of alumina - aluminum phosphate composites fabricated by SLS using different alumina particle sizes and varying amounts of aluminum phosphate in the composite.I--*-- Mean alumina particle size =SO.O microns ....--.. r - - - Mean alumina particle size = 69.0 microns VI--El-- Mean alumina particle size =9.0 micron:xx10.450.500.550.60Relative Density (= I-P)Fig. 2. Variation of the compressive strength of the alumina - aluminumphosphate composites with their relative densities and alumina particle sizes.49In equation 7, the term (1-P) accounts for the dilution in the strength of thecomposite due to porosity. The exponent of the (1-P) term was assumed to be 1 based on the experimental observation in Fig. 2. Vb accounts for the area of the aluminum phosphate through which the crack has to propagate. The a-0.5 term is introduced to characterize the influence of the surface crack size on the material strength.

The a-0.5 term is introduced to characterize the influence of the surface crack size on the material strength. Theexponent of the crack length is -0.5 for mixed mode fracture that occurs duringcompressive loading [10]. Band p are constants. Assuming that the material is completely dense (P=O) and Vb=1, Eqn. 7 reduces to ---- (8)0' =B*a-0.5From Eqn. 8, it can be inferred that B should correspond to the fracture toughness of apure aluminum phosphate specimen with no pores under compressive loading. No datafor the fracture toughness of aluminum phosphate is available in the literature. However, the mode I (KIC) fracture toughness of alumina is between 3 - 5 MPa~m [10].As a first approximation, assume that for the coarsest alumina (80Jlm) grain size equals the crack size. By plotting In(Vb) vs In«j*~a/(1-P)), B and p were determined to be 5.0MPa/~m and 2.6 respectively. These values were then substituted in Eqn.

By plotting In(Vb) vs In«j*~a/(1-P)), B and p were determined to be 5.0MPa/~m and 2.6 respectively. These values were then substituted in Eqn. 7 to calculate the strength of the composites containing 69.0Jlm and 9.0Jlffi aluminapowder assuming that the crack size in each composite equals the alumina grain size(Figs. 3 and 4).Comparing the actual and calculated strengths of the alumina-aluminumphosphate composites in Figs 3 and 4,it can be inferred that equation 7 predicts thetrends in the strength behavior of the composites accurately. However, the calculatedvalues of the strength of the composite assume that the flaw size equals the averagealumina grain size. But as seen in Table 1, there is a size distribution for the alumina particles. Hence the intrinsic crack size is greater than the average alumina particle size.This results in an overestimation of the composite strength if we assume that the grain size equals flaw size in the material (Figs. 3 and 4).

3 and 4). Processing flaws as large as 400Jlmhave been observed on surface of composites made with fine alumina powder (9Jlm). If aflaw size of 400Jlm is used in Fig. 4, then the agreement between calculated andmeasured values improves. Further microstructural analysis is essential to understand theeffect of material and processing parameters on the flaw size in the material.50Alumina grain size =69.0 J.1mMeasured strength (MPa) -0 - - Calculated strength (MPa)0.200.250.300.350.400.45Binder volume fraction (V.)Fig. 3. Measured and calculated compressive strength of alumina .. aluminumphosphate composites fabricated by SLS. Eqn. 7 was used in calculating the strength. It was assumed that the crack size = alumina particle size = 69Jlm, B = 5.0MPa"m and p = 2.6.18Ur------------------,Measured strength (curve a)..... -0..... Calculated strengthp,Defect size=9.Ollffi (curve b)...•../::s.....

-0..... Calculated strengthp,Defect size=9.Ollffi (curve b)...•../::s..... Calculated strengthDefect Size=400J.1ffi (curve cJ..'··b..'.'c ......................······A·······.0.350.5 Binder volume fraction (Vb)0.40.450.55Fig. 4. Calculated and measured compressive strength of alumina .. aluminum phosphate ceramic composites fabricated by SLS. The calculated values were determined51using B=5.0MPa"m and p=2.6 in Eqn. 7. Curve b assumes a crack size of 9.0 flm (alumina particle size in initial blend) and curve c assumes a crack size of 400 Jlffi.CONCLUSIONSThe compressive strength of the alumina - aluminum phosphate composites fabricated by SLS was studied as a function of the porosity, binder contentand the surface flaw size. The strength of the composite increases as the binder contentincreases and as the porosity and defect size decrease.

The strength of the composite increases as the binder contentincreases and as the porosity and defect size decrease. Hence factors such as aluminaparticle size and binder content which influence the microstructure have a strong effecton the strength of the composites. A constitutive equation was formulated to relate theeffects of the microstructural factors on the strength of the composite. While the equation predicted the general trends in the variation of the strength with microstructural factors,more detailed investigation into the crack size controlling the strength of the material isrequired in order to use the constitutive equation as a predictive tool.REFERENCES 1. Proceedings of the Solid Freeform Fabrication Symposium, Aug. 11-14, 1991,University of Texas, Austin, Eds. J.J.Beaman, H.L.Marcus, D.L.Bourell and IW.Barlow.2. C.R.Deckard and J.J.Beaman, "Solid Freeform Fabrication and Selective Laser Sintering", Proc.

viewed PaperEXPERIMENTAL STUDY OF WING STRUCTURE GEOMETRY TO MITIGATE PROCESS-INDUCED DEFORMATIONJustin D. Valenti∗‡, Joseph Bartolai†, and Michael A. Yukish∗ ∗Department of Aerospace Engineering †Applied Research Laboratory The Pennsylvania State University, University Park, PAAbstractSmall uncrewed aerial vehicles that are fabricated with material extrusion additive manu- facturing often have wings that are single-perimeter structures with sparse internal structure. The large distance between internal supports creates an “unsupported-wall distance”, which leaves the wing skin prone to deformation during fabrication. This work explores and quan- tifies the relationship between the deformation of the wing skin and three geometric param- eters: (1) unsupported-wall distance, (2) local surface curvature, and (3) extrusion width.

A three-level full factorial study was devised in which wing sections of varying surface curva- ture, unsupported-wall distance, and extrusion width were fabricated with polymer material extrusion additive manufacturing. The surfaces of the wing sections were then digitized into point clouds with a coordinate measuring machine, and the point cloud data were directly compared to the GCode used to print each wing section. The deformation data was analyzed to quantify the relationship between deformation and the experimental parameters. From the experiment, a non-dimensional term was identified that captures a bounding relationship between the geometric parameters and the deformation. Finally, a mathematical expression was developed to serve an upper bound on unsupported-wall distance based on extrusion width and surface curvature.IntroductionAdditive manufacturing (AM) has emerged as a valuable new fabrication method for small uncrewed aerial vehicles (UAVs).

The design freedom provided by AM has both simplified UAV fabrication and made new families of structure practical to fabricate. While several polymer AM processes have been used to build UAVs or UAV parts [1–11], the low-price point and ubiquity of polymer material extrusion additive manufacturing (MEAM) has lead to many UAVs that were designed and built with MEAM [4, 12–27].The high strength-to-weight requirement of wing structures, combined with the minimum feature size of most desktop 3D printers, typically leads to MEAM wings that have a single perimeter, with sparse internal structure. Additive manufacturing is known to struggle with this type of structure.Previous work by the authors developed a structural design method that produced wing structures for small MEAM-built wings [28].

While wings designed with this method were successfully flown, the main drawback was that the single-extrusion perimeters on the result- ing wings would deform during fabrication due to lack of support. A study was conducted to‡Corresponding Author: jdv5076@psu.edu1704explore the tradespace of weight, aerodynamic performance, and structural performance [29].

The main conclusions of this study were:Undeformed wing skin provides significantly more bending inertia than required(for a standard desktop 3D printer with a 0.4mm nozzle diameter and the chosen sUAV: span = 4ft, weight = 1 lb).Spar spacing and skin curvature strongly affect skin deformation.Additional internal structure reduced deformation to a point where aerodynamic per- formance was minimally affected, with negligibly more additional weight than the baseline internal structure designed to only carry the flight loads.The present work is a continuation of that initial study and aims to further characterize the relationship of geometric design parameters to wing skin deformation.Novel Contribution: The results of the experimental study show a clear bounding rela- tionship between deformation of the single extrusion perimeter and (1) unsupported-wall dis- tance, (2) surface curvature, and (3) extrusion width.

A non-dimensional “unsupported-wall index” is identified which captures a bounding relationship between the geometric param- eters and deformation. Lastly, a mathematical expression is presented to serve as a design guideline for unsupported-wall distance based on curvature and extrusion width.Layout of Paper : The following section provides the methods used to design the wing sec- tions, design the experiment, and obtain and process the deformation data. This is followed by the Results section, which shows the absolute deformation results, non-dimensionalized deformation results which match a previous study, and the presentation of a single non- dimensional value which appears to characterize the problem. Then, a design constraint for thin-walled structures presented based on the study data and analysis. Finally a conclusion section offers a small discussion of the results, concluding remarks, and future work.

Finally a conclusion section offers a small discussion of the results, concluding remarks, and future work. A nomenclature section is also provided at the end of the manuscript.MethodsWing Design and GCode Generation: Custom software, named GWing, was developed to simultaneously design UAV wings and generate the GCode to build them. GWing was used to design the wing sections and generate the GCode used for this study. The software breaks wings into printable wing sections. Each wing section is fabricated with spanwise direction of the wing aligned with the build direction and the chordwise direction aligned with the X-axis of the build plate, as shown in Figure 1. A brim is used to promote adhesion to the build plate. A solid first layer is also used which both improves build plate adhesion and becomes a rib in the final wing structure. All subsequent layers consist of a single perimeter and the internal structure.

All subsequent layers consist of a single perimeter and the internal structure. In each layer, the single perimeter is printed first which corresponds to the desired airfoil. The internal structure is printed second and consists of one or more vertical members, which become spars in the final wing structure. The tool path for an arbitrary layer is shown in Figure 2(a).Experimental Design: A three-level full factorial study was devised to test the effects of surface curvature, κ, spar spacing, dx, and extrusion width, we, on the deformation of the perimeters during the fabrication process. A parameterized wing section is shown in1705Figure 1 : Example GCode of wing section, annotated with chord and section span dimensions. Extrusion moves are in red. Travel moves are shown in grey.(a) Toolpath for an arbitrary layer(b) Section ParameterizationFigure 2 : (a) Illustrates the toolpath for a layer of a wing section.

Extrusion moves are in red. Travel moves are shown in grey.(a) Toolpath for an arbitrary layer(b) Section ParameterizationFigure 2 : (a) Illustrates the toolpath for a layer of a wing section. The outer perimeter is extruded first, followed by the internal structure from the trailing edge to the leading edge. (b) Diagrams the parameterization of a wing section. Chord, c, was set to 122 mm for all wing sections in this study. Spar spacing, dx, curvature, κ, and extrusion width, we, were the varied parameters in this study.Figure 2(b). The wing shape and structure was constant for every layer. The chord for every section was 122mm.Both the location and number of spars were defined by a single parameter, dx, the spanwise distance between spars. The high aspect ratio of an airfoil section means that spar spacing is roughly equal to unsupported-wall distance, dx ≈ du. It is assumed that any effect on deformation from du with also appear with dx.

It is assumed that any effect on deformation from du with also appear with dx. The three levels of spar spacing were dx = {20, 50, 80}mm.Curvature is defined as the inverse of radius of curvature, κ = 1/ρ. For a given functionz(x), curvature is calculatedκ(x) =|z′′| (1 + z′2)(3/2) .Note, the dimension of curvature is length−1. While the curvature of a typical airfoil varies across the chord for both the upper and lower surfaces, airfoil sections for this experiment consist of circular arcs on both the upper and lower surface, as shown in Figure 2. The curvatures of both the upper and lower surfaces were equal. The three values of curvature1706(1)used were κ = {0, 2.5, 5}m−1.To maintain the proportions of nozzle diameter, extrusion width, and layer height across all test sections, three MEAM machines were used, each with a different nozzle diameter.

A Lulzbot Mini 2 with a nozzle diameter of 0.25mm was used to fabricate sections with a design extrusion width we,d = 0.31mm and a layer height hl = 0.17mm. A Creality CR-10 with a nozzle diameter of 0.40mm was used to fabricate sections with a design extrusion width we,d = 0.50mm and a layer height hl = 0.27mm. A second Lulzbot Mini 2 with a nozzle diameter of 0.50mm was used to fabricate sections with a design extrusion width we,d = 0.62mm and a layer height hl = 0.33mm.In summary, three parameters were varied at three levels to provide 27 test points, withthe parameter levels being80 }mm, 5 }m−1, we,d = { 0.31, 0.50, 0.62 }mm.dx = { κ = {20, 0,50, 2.5,Sectional views of the variation in curvature and spar spacing are shown in Figure 3. A wing section was designed and fabricated for each point. The deformation of both the upper and lower surfaces were independently quantified, doubling the number of data points.

A wing section was designed and fabricated for each point. The deformation of both the upper and lower surfaces were independently quantified, doubling the number of data points. The build parameters were:MEAM Machine∗: { Lulzbot Mini 2, Creality CR-10, Lulzbot Mini 2† } • Slicing Software: GWing (custom, described above) • Material: • Nozzle Diameter∗: { 0.25, 0.40, 0.50 }mm • Layer Height∗: { 0.17, 0.27, 0.33 }mm • Nozzle Temp: 210 deg C • Bed Temp: 60 deg C • Toolhead Speed: 40 mm/sPolylactic Acid (Essentium PLA XTR Natural)For the purposes of this study, differences between the three MEAM machines are assumed to be negligible.Measurement: Calipers were used to measure the wall thickness of each section for both the upper and lower surfaces. This provided a “measured extrusion width”, we,m, to com- pare to the designed extrusion width, we,d.

This provided a “measured extrusion width”, we,m, to com- pare to the designed extrusion width, we,d. A laser coordinate measuring machine (CMM), specifically a FARO arm [30], was used to digitize the surface of each wing section in a point cloud. A wing section on the CMM table, along with the point cloud capturing its surface are shown in Figure 4.Each point cloud was directly compared to the corresponding GCode file. For each point in a point cloud, a local deformation, δ, was calculated. The root mean square (RMS) of all deformation values for a particular section was calculated to provide an RMS deformation of the as-built part. This process is provided in Algorithm 1.∗Parameters listed in sets of three correspond, respectively, with the three we,d values. †Two separate Lulzbot Mini 2s were used.

This process is provided in Algorithm 1.∗Parameters listed in sets of three correspond, respectively, with the three we,d values. †Two separate Lulzbot Mini 2s were used. One with a nozzle diameter of 0.25 mm and the other with anozzle diameter of 0.5 mm.1707Figure 3 : Three-level factorial design for unsupported-wall distance, dx, and curvature, κ. The 9 design points illustrated above were each repeated 3 times for extrusion widths we,d = [0.31, 0.50, 0.62]mm. For all wing sections, the chord c = 122mm.Algorithm 1: Compare a point cloud, defined by a list of points P = (p1, p2, . . . , pm), to a GCode file to calculate deformation of external surface of as-built part.Input: GCode, Point Cloud Output: RMS Deformation of As-Built Part1 Convert Extrusion moves in GCode to list of points G = (g1, g2, . ..

. . , gn); 2 Align GCode points and point cloud; 3 for each GCode Layer do 4Find all Point Cloud point within half a layers height; for Each Point Cloud Point, pi, at the current layer do Find the closest GCode point at that layer, gj to pi; Determine whether gj+1 or gj−1 is the next closest to pi; Calculate the local deformation, δi, defined as the perpendicular distance between pi and the line defined by the two closest GCode points;5678end9 10 end 11 Calculate the RMS of all δ values to find to RMS deformation of the as-built part ;The RMS deformation provided the measure of a accuracy for a particular wing section and was used to quantify the deformation that occurred during the build process. The relationship between deformation and the design parameters is explored in the next section.ResultsDimensional Results: First, the RMS deformation data, in millimeters, is plotted against the dimensional design parameters in Figure 5.

To allow for the use of symbols to convey data, design extrusion width is used in Figures 5(a) and 5(b). Measured extrusion width is used in Figure 5(c). Clear bounding behaviors are shown in all three plots in Figure 5. Qual-1708(a) Wing Section On Table(b) Point Cloud(c) Point cloud colored by local deformationFigure 4 : Example wing section from GCode illustrated in Figure 1. (a) The fabricated wing section on mounted on the CMM table awaiting measurement. (b) The point cloud generated by the CMM. (c) The point cloud with the deformation of each point from the intended GCode calculated.itatively, it can be observed that as expected, RMS deformation increases with increased unsupported-wall distance, decreases with increased curvature, and decreases with increased extrusion width. Analysis of Variance (ANOVA) was performed to determine if the effects of spar spacing, curvature, and extrusion width on RMS deformation were statistically signif- icant.

Analysis of Variance (ANOVA) was performed to determine if the effects of spar spacing, curvature, and extrusion width on RMS deformation were statistically signif- icant. The p-values of the three experimental parameters, as well as the interaction terms, are provided in Table 1. The small p-values indicate that the null hypothesis can be rejected. This suggests that all parameters and their interactions have a statistically significant affect on deformation.Table 1: P-values of the three experimental parameters and their interaction terms.Parameter P-valuedx 8.28e-13κ 6.04e-12we,d 1.79e-9dx × κ κ × we,d we,d × dx 7.80e-6 4.89e-6 2.87e-6Non-Dimensional Results: In an effort to combine design parameters and simplify theproblem, RMS deformation, unsupported-wall distance, and curvature were all non-dimensionalized by measured extrusion width. Both RMS deformation and unsupported-wall distance were divided by extrusion width.

Both RMS deformation and unsupported-wall distance were divided by extrusion width. Curvature was multiplied by extrusion width, since the di- mension of curvature is length−1. Similar to the dimensional results, non-dimensional RMS deformation generally increases with non-dimensional unsupported-wall distance and gener- ally decreases with non-dimensional curvature, as shown in Figure 6.In an effort to further collapse the problem to a single dimensionless parameter, theunsupported-wall index, η, was conceived and is defined asη =(cid:18) κ dx(cid:19)w2 e.Non-dimensional RMS deformation is plotted versus η in Figure 7. As with the rest of this work, plotting deformation versus unsupported-wall index does not show a clean curve, but1709(2)(a) RMS deformation vs. spar spacing, colored by curvature, symbols representing design extrusion width(b) RMS deformation vs.

spar spacing, colored by curvature, symbols representing design extrusion width(b) RMS deformation vs. curvature, colored by spar spacing, symbols representing design extrusion width(c) RMS deformation vs. measured extrusion width, colored by spar spacing, symbols representing curvatureFigure 5 : RMS deformation plotted versus the three design parameters. These plots show that, in general, deformation increases with spar spacing, decreases with curvature, and decreases with extrusion width.(a) RMS deformation vs spar spacing, colored by curvature, all three values non-dimensionalized by measured extrusion width(b) RMS deformation vs curvature, colored by spar spacing, all three values non-dimensionalized by measured extrusion widthFigure 6 : Deformation, spar spacing, and curvature non-dimensionalized by measured extrusion width. This allows the 3 dimensional design parameter to be collapsed to two non-dimensional parameters.

Rearranging Equation 2 yieldsdu <1 η∗ κw2ewhere η∗ is a critical unsupported-wall index. As noted in the previous section, the present study suggests 2 × 10−5 as reasonable value. Unfortunately, equation 3 drives unsupported- wall distance to zero as curvature goes to zero. Plotting non-dimensional deformation versus non-dimensional spar spacing, as shown in Figure 6(a), suggests that constraining the ratio of spar spacing to extrusion width to less than 50 yields an RMS deformation of less than one extrusion width. Translating this into a constraint on unsupported-wall distance yieldsdu < r∗we,where r∗ is a critical ratio of unsupported-wall distance to extrusion width. Combining the two constraints yields(cid:20) 1 η∗ κw2 Equation 5 provides a constraint on unsupported-wall distance based on curvature and ex- trusion width. η∗ and r∗ are tuning parameters.

Combining the two constraints yields(cid:20) 1 η∗ κw2 Equation 5 provides a constraint on unsupported-wall distance based on curvature and ex- trusion width. η∗ and r∗ are tuning parameters. Suggested values based on the present data are [η∗, r∗] = [2 × 10−5, 50].(cid:21)e , r∗wedu < max.ConclusionAn experimental study was conducted to determine the effect of unsupported-wall dis- tance, curvature, and extrusion width on the deformation of the skin of MEAM-produced wing sections during the fabrication process. While a precise mathematical function of the1711(3)(4)(5)design parameters for deformation could not be identified, the deformation data exhibited clear bounding behavior. The data showed that, in general, deformation increased with unsupported-wall distance, decreased with curvature, and decreased with extrusion width. Similar trends were identified when non-dimensionalizing deformation, unsupported-wall dis- tance, and curvature by extrusion width.

Similar trends were identified when non-dimensionalizing deformation, unsupported-wall dis- tance, and curvature by extrusion width. The results of an ANOVA study suggested that these trends were statistically significant.Further non-dimensional analysis identified a dimensionless “unsupported-wall index”, η, which appears to collapse the problem to a single parameter. The wall deformation appears to be relatively small at higher η values.Finally, a novel constraint on unsupported-wall distance was developed and presented in Equation 5. The expression constrains unsupported-wall distance based on curvature and extrusion width.Future work will apply the new design constraint to different problems and explore the validity of the new design constraint. Additionally, there is work to be done to explore the physical mechanisms of deformation that were observed in the this study.AcknowledgementsThe authors would like to offer special thanks to Dr.

Additionally, there is work to be done to explore the physical mechanisms of deformation that were observed in the this study.AcknowledgementsThe authors would like to offer special thanks to Dr. Charlie Tricou, for his expertise in design of experiments and helping to conceive the origins of this work, and Dr. Callie Zawaski and Dr. Simon Miller for their many hours of discussion on this work and how to improve it.References[1] S. K. Moon, Y. E. Tan, J. Hwang, and Y. J. Yoon, “Application of 3D printing technol- ogy for designing light-weight unmanned aerial vehicle wing structures,” International Journal of Precision Engineering and Manufacturing - Green Technology, 2014.[2] S. J. Furst, G. Bunget, and S. Seelecke, “Design and fabrication of a bat-inspired flapping-flight platform using shape memory alloy muscles and joints,” Smart Mate- rials and Structures, vol. 22, no. 1, 1 2013.[3] C. Richter and H.

PERIMENTAL TEMPERATURE ANALYSIS OF POWDER-BASED ELECTRON BEAM ADDITIVE MANUFACTURINGSteven Price*, James Lydon**, Ken Cooper**, Kevin Chou*Department of Mechanical Engineering The University of Alabama, Tuscaloosa, AL 35487, USA**Advanced Manufacturing Technology Team Marshall Space Flight Center, Huntsville, AL 35812, USAAbstractA near infrared thermal imager has been employed for temperature measurements (build part surfaces) in the powder-based electron beam additive manufacturing (EBAM) process. A methodology has also been developed to analyze temperature distributions and history around the melting scan area. The temperature profiles along the beam scanning clearly show the moving source of heat phenomenon, with the peak temperature reaching over 2000 °C (for Ti- 6Al-4V) followed by extremely rapid cooling opposite to the beam scanning direction, except a slow-cooling portion corresponding to the liquidus-solidus range.

The build surface temperatures and the melt pool sizes, in the localized electron beam scanning area, were studied at various configurations, e.g., different build heights, also with or without an overhang.IntroductionAdditive manufacturing (AM) is a fabrication technique that creates parts layer by layer based on geometry defined by computer-aided design (CAD) software. Many different methodologies of AM exist such as fused deposition modeling, stereolithography, and selective laser sintering. Some types produce plastic parts while others are capable of fabricating metal parts. Several benefits of AM are rapid production time, capability to produce complex geometry, and no tooling (Gong et al., 2012) [1].Electron beam additive manufacturing is an AM technology that uses an electron beam to melt cross-sections of a part layer after layer in a metal powder such as Ti-6Al-4V.

EBAM is capable of producing high strength, full density metal parts unlike other AM methods such as selective laser sintering. Because of its high energy density, the EBAM fabrication process is much faster than alternatives such as direct laser melting. EBAM enables the fabrication of metal parts with complex geometries such as fine network structures and internal channels and cavities that are difficult to produce using conventional manufacturing methods (Gong et al., 2012) [1].The EBAM build process consists of pre-heating, contour melting, and hatch melting phases. The pre-heating phase elevates the temperature of the powder in order to reduce the thermal gradient experienced during melting, thereby reducing residual stresses. Contour melting is accomplished by multiple electron beams that melt the boundary of the cross section being created.

Contour melting is accomplished by multiple electron beams that melt the boundary of the cross section being created. Hatch melting is performed by a single electron beam scanning a raster pattern inorder to melt the powder that will become the interior of the part. A new layer of powder is then distributed over the build platform in preparation for the next layer’s fabrication.The ability to measure the temperatures that occur during the build process is an essential capability that must be established in order to verify process models, predict part microstructure, and develop feedback control systems. Thermocouples and many types of infrared cameras have been used to measure process temperatures in various types of AM processes. Zäh and Lutzmann (2012) [2] used a thermocouple to measure the temperature 0.3 mm below the powder layer during an EBAM process. Qian et al.

Zäh and Lutzmann (2012) [2] used a thermocouple to measure the temperature 0.3 mm below the powder layer during an EBAM process. Qian et al. (2005) [3] used four Type K thermocouples and three Type W thermocouples to monitor temperatures in the substrate and the part interior during material deposition during a direct laser fabrication (DLF) process. Thermocouples, however, tend to have several limitations in their effectiveness at monitoring process temperatures in additive manufacturing. Because additive manufacturing involves a fast, transient heat source, the limited spatial temperature resolution and slow response time of thermocouples limit their usefulness when monitoring AM processes.Non-contact infrared imagers have been widely utilized in AM temperature measurement because of their excellent spatial temperature resolution and fast response time.

Wegner and Witt (2011) [4] used an InfraTec Jade III MWIR thermal imaging system with a spectral range of 3-5 µm, a resolution of 320 x 240 pixels, and a frame rate of 700 Hz to measure the temperature distribution during a laser sintering process. Infrared cameras sensitive in the near-infrared spectrum (0.8-2.5 µm) are better suited for measuring high temperatures (>1000 K) and also provide the benefit of being less sensitive to surface emissivity. Craeghs et al. (2011) [5] used a high-speed NIR CMOS camera and a photodiode sensor that were both sensitive to the spectral range of 0.7-0.95 µm to monitor the size of the molten pool during a layer-wise laser melting process. Wang et al.

Wang et al. (2009) [6] used a Thermaviz two-wavelength imaging pyrometer system sensitive in the 0.7-0.8 µm and 0.8-0.9 µm wavelengths and a detectable temperature range of 1450-1860 °C to measure the temperatures and molten pool shape during a LENS process.Recently, there have been a few attempts for the use of infrared cameras to measure temperatures during the EBAM process. Schwerdtfeger et al. (2012) [7] used a FLIR A320 IR camera with a resolution of 320 x 240 pixels and a spatial resolution of 0.83 mm per pixel to detect flaws during an EBAM process. In order to avoid the metallization of the camera’s viewport, a mechanical shutter was used to cover the viewport during the melting sequence. The shutter was then opened immediately after the layer’s melting sequence was completed so the IR camera could capture images. The captured IR images had a low resolution; however the general locations of flaws were detectable. Rodriguez et al.

The captured IR images had a low resolution; however the general locations of flaws were detectable. Rodriguez et al. (2012) [8] used a FLIR SC635 IR camera with a resolution of 640 x 480 pixels and a spatial resolution of 350 µm per pixel as part of a feedback control system for an Arcam A2 EBAM system. They also used a shutter to shield the camera-port during the melting sequence, opening it for viewing once the melting was finished. Non-uniform temperature distributions were easily identified with the IR camera. An issue in determining accurate relative temperatures seen by both Schwerdtfeger et al. [7] and Rodriguez et al. [8] was the apparent “cold spots” that can be caused either by lower temperatures or lower surface emissivity. This is one of many difficulties that must be faced in order to achieve accurate temperature measurements in the EBAM process with an infrared camera.

This is one of many difficulties that must be faced in order to achieve accurate temperature measurements in the EBAM process with an infrared camera. Radiation from the electrons requires that the viewport into the build area be shielded with leaded glass.This glass causes transmission losses in the infrared spectrum. In order to capture images during the melting process the shutter must be opened which causes a metal film to be deposited on the surface of glass. This film causes further transmission losses. The emissivity of the target area is a critical parameter when converting radiant temperature to true temperature. Emissivity is dependent on surface condition, wavelength, temperature, and angle of incidence (González- Fernández et al., 2012) [9]. Because emissivity is dependent on so many factors, it can be a very difficult property to determine. Price et al.

Because emissivity is dependent on so many factors, it can be a very difficult property to determine. Price et al. (2012) [10] evaluated a near infrared (NIR) thermal imager to measure temperature distributions and history of a build part surface during the EBAM process. It was shown that the NIR thermography, with a shorter spectral range, is able to capture temperature images of the preheating, contour melting, and hatch melting phases of the EBAM fabrication process. Moreover, the molten pool shape and size as well as an average molten pool temperature profile can be determined. A more detailed literature review and explanation of the experimental setup used in this paper may be found in Price et al. [10].In this study, the repeatability of the 2D temperature profiles as well as a more thorough study of molten pool dimensions are presented.

[10].In this study, the repeatability of the 2D temperature profiles as well as a more thorough study of molten pool dimensions are presented. The emissivity of the molten pool and the transmission loss due to metallization were calculated, and the studies of build height and overhang effects were also conducted.Experimental SetupAn Arcam S12 EBAM machine at NASA’s Marshall Space Flight Center was used to build the test parts. Temperature measurements of the EBAM process were made with a LumaSense MCS640 NIR infrared camera. The NIR camera is has a spectral range of 0.78-1.08 µm and a 640 x 480 pixel uncooled focal plane array (FPA) sensor. The image capturing system has a maximum frame rate of 60 Hz and a detectable temperature range of 600 to 3000 °C which is divided into three nominal ranges: low (735-1108 °C), medium (1057-1485 °C), and high (1503-2446 °C). The camera has a lens with a minimal 500 mm working distance and a view area of 31 mm x 23 mm at 500 mm distance.

The camera has a lens with a minimal 500 mm working distance and a view area of 31 mm x 23 mm at 500 mm distance. The camera was calibrated with the glass in the viewport of the S12 EBAM machine so that a transmission rate of unity will correctly account for the transmission losses due to the glass. The powder material used was Ti-6Al-4V. The emissivity of the surface was unknown so measurements were taken with assumed emissivities of 0.35 and 0.5. The integration time of the image capturing system is 16.25 ms, 1.7 ms, and 50 µs for the low, medium, and high temperature ranges, respectively. The data acquisition system consisted of a laptop and an Ethernet cable, connected with a gigabit Ethernet express card.The NIR camera was mounted on a tripod and positioned to look downward through the S12 EBAM machine’s viewport onto the build platform. The camera was vertically inclined at 35° from the surface normal with minimal roll angle and horizontal yaw.

The camera was vertically inclined at 35° from the surface normal with minimal roll angle and horizontal yaw. More detail on the experimental setup may be found in Price et al. (2012) [10]. The S12 EBAM machine contains a heat shield that is placed between the viewport and the build platform. This heat shield has a cutout for viewing the build platform but that cutout is obstructed with vertical bars. Three of these bars were cut to provide a clear view of the build platform, but interference from a single bar was still observed during testing. Several parts of different geometries were fabricated in order to study a variety of effects. Simple rectangular cross-sections were employed in order to study hatch melting and overhang structures were used to study melting over a powder substrate.Results and DiscussionA.

Simple rectangular cross-sections were employed in order to study hatch melting and overhang structures were used to study melting over a powder substrate.Results and DiscussionA. Near Infrared Imaging of the EBAM ProcessThe different temperature ranges of the NIR camera are each uniquely suited for imaging the various stages of the fabrication process. The low temperature range is best suited for preheating, the medium temperature range for contour melting, and the high temperature range for hatch melting. Examples of NIR images of the various portions of the EBAM process are shown in Figure 1. The images are from the fabrication of a 30.48 mm x 25.4 mm solid cross- section.The spatial resolution of the NIR images was determined by identifying the edges of the part during the contour melting phase, then determining the number of pixels between these edges.

The average horizontal and vertical spatial resolutions of the images were found to be 46.8 µm/pixel and 66.2 µm/pixel respectively. The vertical spatial resolution is larger due to distortion caused by the 35° tilt of the camera.Preheating (Low Temperature Range)Contour Melting I (Med. Temperature Range)Contour Melting II (Med. Temperature Range)Hatch Melting (High Temperature Range)Figure 1. NIR images of the various stages of the EBAM fabrication process.B. Repeatability of Temperature MeasurementsThe repeatability of the temperature measurements was also examined. This was accomplished by creating 2D temperature profiles along the scan path of each frame and averaging these profiles together. The average 2D temperature profile of the hatch melting of a single layer with standard deviations is shown below in Figure 2. For this particular layer, theelectron beam scanned back and forth horizontally.

For this particular layer, theelectron beam scanned back and forth horizontally. Temperature profiles of leftward scans were flipped to be rightward and all the profiles were aligned at their peak temperatures to correctly line up the profiles for averaging. The temperature profile appears to be very repeatable, especially in the high temperature region around the beam location.280025002200)C( e r u t a r e p m e T190016001300100015105 Distance (mm)05Figure 2. Average hatch melting temperature profile from a single layer.Average temperature profiles were made for four different layers over a ten layer range (0.7 mm range) and compared together in order to study the repeatability of the temperature measurements over multiple layers. These profiles are shown in Figure 3 and it may be seen that the temperature profiles match almost identically. Layer 17’s temperature profile shows a plateau slightly above 1800 °C where the peak should be found.

Layer 17’s temperature profile shows a plateau slightly above 1800 °C where the peak should be found. This occurred because the footage of this layer was captured using the medium temperature range of the camera. The plateau occurs at that the maximum detectable temperature of the medium temperature range.28002500)C( e r u t a r e p m e T22001900Layer 17 (1.12 mm)Layer 19 (1.26 mm)Layer 25 (1.68 mm)1600Layer 27 (1.82 mm)1300100015105 Distance (mm)05Figure 3. Average profiles showing the repeatability of the temperature measurements.C. Build Height EffectsThe effect of build height on the thermal behavior of the powder during hatch melting was investigated. An average 2D temperature profile along the scanning path was generated for each recorded build height and then compared in Figure 4 to identify trends.

An average 2D temperature profile along the scanning path was generated for each recorded build height and then compared in Figure 4 to identify trends. The horizontal distance from the peak temperature to the solidus temperature on the cooling side of the profile is shorter at higher build heights; however the location of liquidus temperature does not appear to vary with build height. A suspected cause for the change in horizontal location of the solidus temperature is higher beam speeds and currents used by the Arcam S12 machine in the initial layers of the build. These results show that build height may have very little effect on the thermal characteristics of the EBAM process.28002500Layer 19 (1.26 mm))C( e r u t a r e p m e T220019001600Layer 25 (1.68 mm)Layer 71 (4.90 mm)Layer 187 (11.02 mm)Layer 299 (20.86 mm)1300100015105 Distance (mm)05Figure 4. Temperature profiles across the molten pool at various build heights.D.

Temperature profiles across the molten pool at various build heights.D. Transmission Loss due to Metallization of Sacrificial GlassDuring the fabrication process, some of the Ti-6Al-4V powder gets evaporated and forms a thin layer on the interior of the viewport. A piece of sacrificial glass shields the thick, protective glass in order to protect it from the metallization. This glass can be easily discarded and replaced after it becomes obscured by the metal film. The sacrificial glasses from several of the experiments were salvaged and tested in order to determine the severity of the transmission loss caused by the metallization. The sacrificial glass used in the experiment that generated all the data for this paper except the overhang study is shown below in Figure 5. The transmission losses were determined by the following experimental procedure. The NIR camera was used to image a tungsten filament with a clear sacrificial glass between the two.

The transmission losses were determined by the following experimental procedure. The NIR camera was used to image a tungsten filament with a clear sacrificial glass between the two. The average temperature of the filament was then determined. Next, the darkened sacrificial glass from the experiment was placed in between the camera and the filament and the average measured temperature of the filament was determined for multiple transmission rates. The true transmission rate was found by selecting the one that provided the filament temperature closest to the one measured with the clear glass. The transmission loss from the metallization of theglass alone in air shown in Figure 5 was found to be about 20% and the transmission loss due to the leaded and vacuum glasses has been measured to be 47%. The total transmission loss caused by the metallization is about 9.4% and this was found by multiplying the two transmission losses together.Viewing AreaFigure 5.

The total transmission loss caused by the metallization is about 9.4% and this was found by multiplying the two transmission losses together.Viewing AreaFigure 5. Sacrificial glass with metallization.E. Molten Pool EmissivityThe emissivity of the molten pools was estimated by determining the measured value of the liquidus temperature on the 2D temperature profile then calculating the emissivity required to covert the measured liquidus temperature to Ti-6Al-4V’s true liquidus temperature (1655 °C).

The relative sensitivity of the camera at λ was assumed to be unity because this information was not available from the camera manufacturer. The equation was numerically solved for a range of apparent temperatures. The apparent temperature that resulted in the smallest value of M was the correct value. Using the calculated temperature as the apparent temperature and the liquidus temperature of Ti-6Al-4V for the true temperature, the equation was numerically solved again, this time looking for the emissivity that resulted in the smallest value of M.The average molten pool emissivities and their standard deviations were calculated at various build heights and are shown in Figure 6. The grand mean is represented by the solid red line and the standard deviation of the averages is represented by the dashed red lines. The grand mean was found to equal 0.283 and the standard deviation 0.018. Boivineau et al.

The grand mean was found to equal 0.283 and the standard deviation 0.018. Boivineau et al. (2006) [12] found the emissivity of Ti-6Al-4V to be around 0.58 at the liquidus temperature where it then drops to 0.4 in its liquid state. The emissivity they found was the normal spectral emissivity of Ti-6Al-4V at a spectral wavelength of 684.5 nm. The NIR camera employed in the current study was sensitive in the 780 nm to 1080 nm range, this might explain the difference in calculated emissivities since emissivity is a wavelength dependent property.0.40.3750.350.325y t i v i s s i m E0.30.2750.250.2250.20510 15 Build Height (mm)2025Figure 6. Average molten pool emissivities with standard deviations at various build heights.F. Molten Pool DimensionsThe average length and width of the molten pool were also determined at various build heights. These dimensions were found by generating and measuring an isotherm at the measured liquidus temperature.

These dimensions were found by generating and measuring an isotherm at the measured liquidus temperature. An example post-processed image of a molten pool during hatch meltingis shown in Figure 7. The temperatures of the white pixels are equal to or greater than the liquidus temperature. The average molten pool dimensions are shown in Figure 8. The grand means of the molten pool average lengths and widths were 2.73 mm with a standard deviation of 0.16 mm and 0.98 mm with a standard deviation of 0.13 mm, respectively.2.01 mm0.86 mmFigure 7. Post-processed NIR image showing size and shape of molten pool.43.53)m m2.5( e c n a t s i D21.5Avg Length Avg Width10.5005101520253035Build Height (mm)Figure 8. Average molten pool dimensions with standard deviations at various build heights.G.

Average molten pool dimensions with standard deviations at various build heights.G. Overhang Structure Thermal EffectsThe thermal effects of melting over a powder substrate were also investigated in order to better understand this common scenario. A powder substrate is encountered during the build process because the powder serves as support material under overhang features of the part. The images shown in Figure 10 are from the fabrication of the middle “shelf” of the part shown in Figure 9. The top-right quarter of each image in Figure 10 has a solid substrate and the bottom-right quarter of each image has a powder substrate. The images from the first overhang layer show that cooling occurs much slower over the powder substrate when compared with the solid substrate. However, this effect is almost completely diminished by the third layer above the powder substrate.

However, this effect is almost completely diminished by the third layer above the powder substrate. In order to study the slower cooling over the powder substrate 2D temperature profiles were generated from frames from the first overhang layer footage. The profiles were aligned at the interface of the solid and powder substrates and oriented so that the common scanning direction was rightward, then averaged together. 2D temperature profiles were generated parallel to the scanning path at intervals of 2 mm out to 8 mm in order to show the cooling effect. The left image in Figure 11 shows an example of the locations of these temperature profiles and the right image shows the average temperature profiles. At 8 mm from the scan path, the part has almost cooled to the same temperature as the surrounding area. The dashed red boxes represent the regions that have a solid substrate.Figure 9. Overhang model and part fabricated during EBAM testing.3rd Overhang Layer Figure 10.

The dashed red boxes represent the regions that have a solid substrate.Figure 9. Overhang model and part fabricated during EBAM testing.3rd Overhang Layer Figure 10. Temperature maps showing hatch melting over a half powder-half solid substrate.1st Overhang Layer11th Overhang Layer12008 mm6 mm4 mm)C( e r u t a r e p m e T11001000900800Scan Direction0 mm2 mm4 mm6 mm2 mm7008 mm0 mm03 9 Distance (mm)612 15Figure 11. Average temperature profiles from first overhang layer.ConclusionsAn NIR camera has been used to make temperature measurements during an EBAM process. A part was built by the EBAM process that was designed to enable the observation of electron beam melting over a powder substrate. The effects build height has on various thermal characteristics of the electron beam melting process have been evaluated. The emissivity and dimensions of the molten pool were also determined at various build heights.

Major conclusions from this research are listed below.2D temperature profiles show that temperature measurements of the EBAM process with an NIR camera are repeatable.Build height may have very little effect on the thermal characteristics of the EBAM process, though some effect on the length of the liquidus-solidus transition region.Metallization of the viewport during fabrication leads to increased transmission losses through the viewport.The calculation of the molten pool emissivity is very sensitive to the measured liquidus temperature.The melting of the first layer over a powder substrate shows a significantly decreased cooling rate when compared to cooling over a solid substrate.Future work will compare experimental data with simulation results in order to validate process models as well as to study process parameter effects on the thermal characteristics of the EBAM process.AcknowledgementThis research is supported by NASA, No.

 EXPERIMENTAL-NUMERICAL INVESTIGATION OF HEAT TRANSFER DURING SELECTIVE LASER MELTINGM. Masoomi1, S. M. Thompson1†, N. Shamsaei1, A. Elwany2, L. Bian31Department of Mechanical Engineering, Mississippi State University, Mississippi State, MS 39762 1Center for Advanced Vehicular Systems (CAVS), Mississippi State University, Mississippi State, MS 39762 2Department of Industrial and Systems Engineering, Texas A & M University, College Station, TX 3Department of Industrial and Systems Engineering, Mississippi State University, Mississippi State, MS 39762†Corresponding author: Email: thompson@me.msstate.edu Phone: (662) 325 2364AbstractThe heat transfer in and around a part being fabricated via Selective Laser Melting (SLM) is numerically simulated while considering the surrounding powder bed modeled to have an effective thermal conductivity. By accurately simulating the powder bed heat transfer during SLM, mechanical properties of parts can be better predicted.

By accurately simulating the powder bed heat transfer during SLM, mechanical properties of parts can be better predicted. Heat transfer to previously- deposited layers and the build plate are also simulated. In order to validate the presented model, a thermocouple was embedded into a substrate used and a SLM system was utilized for performing two experiments. In the first set, various laser power and scan speed combinations were employed while passing the laser over the thermocouple-embedded substrate. This procedure calibrated the numerical model and demonstrated that the heat transfer due to convection and radiation during deposition of a single layer is approximately 10-15% of initial laser power input. The final experiment consisted of building a thin wall of SLM of 17-4 PH stainless steel (SS). The effects of scan pattern and part size on the temperature response of and around the part are demonstrated as significant.

The effects of scan pattern and part size on the temperature response of and around the part are demonstrated as significant. Distinct heating and cooling rates are also provided for these various cases; indicating the dependency of final microstructure on part size and the utilized scan pattern.IntroductionSelective Laser Melting (SLM), a Powder Bed Fusion (PBF) additive manufacturing (AM) method, is a unique means for producing complex parts not possible via traditional manufacturing methods. During SLM, a laser beam melts powder material layer by layer to produce a part via solidification. Within an inert atmosphere, a micro-sized melt pool (i.e. molten pool) is induced by a large heat flux via irradiation from a relatively small-diameter laser beam.The melt pool moves along with the laser beam and solidifies very rapidly upon laser removal, forming a track of solid material.

This process is repeated to form multiple tracks with an imposed raster/scan strategy until all individual layers are deposited. After finishing the build, the structure is detached from the substrate, typically via Electrical Discharge Machining (EDM), and then excess powder is removed. SLM methods offer several advantages over traditional ‘subtractive’ manufacturing methods. Parts with complex geometries can be built more easily and small-lot, customized-application production lines can become more efficient.Due to the focused laser beam inducing high heat flux via irradiation, SLM parts experience very high, and localized, temperature gradients and heating/cooling rates, resulting in relatively fast melting and solidification. The solidification and cooling rates directly affect and set the microstructure of built parts.

The solidification and cooling rates directly affect and set the microstructure of built parts. Knowledge of the solidifying melt pool behavior upon removal of the laser, including solid phase nucleation and dendritic growth, will indicate the initial microstructure phase and distribution. Secondary solidification, or microstructural evolution, continues at a longer time scale during SLM due to solid phase transformations resulting from cyclic laser-to-substrate conduction. These single and multi-phase heat transfer phenomena are very sensitive to SLM process parameters, such as laser power, laser beam size and traverse speed. In order for SLM parts to function reliably in engineering applications, the relationship between microstructure and final part mechanical properties is paramount [1]. This relationship can be determined either experimentally, numerically or analytically.

This relationship can be determined either experimentally, numerically or analytically. Since experimental, trial-and-error SLM research can be costly and time-consuming, there is an appeal to use validated numerical simulations for learning process-parameter relationships.Many have simulated the SLM fabrication of various materials with varying degrees of fidelity in the past several years [2, 3]. Performing numerical simulations allows for time/energy/materials savings by learning process-parameter relationships without repetitious trial-and-error experimentation. In order to numerically simulate powder melting and melt pool solidification during SLM, latent thermal energy transfer needs to be idealized and/or rigorously incorporated into the model. In addition, the simulation of the transient temperature distribution in and around the part during manufacturing needs to be predicted in order to estimate the encumbered part microstructure.

In addition, the simulation of the transient temperature distribution in and around the part during manufacturing needs to be predicted in order to estimate the encumbered part microstructure. The SLM process parameters play a significant role in governing thermal history and thus microstructure type, size and orientation within the part. Via numerical simulation, the sensitivity of thermal history, and thus final microstructure, on various part shapes/sizes and various operating environments (e.g. radiation and convection heat transfer) can be readily performed. In addition, guidelines for effective SLM process control can be ascertained by exploiting any learned relationships between microstructure and heat transfer.In order to generate trustworthy numerical models, thermal diagnostics of the SLM process is needed in order to benchmark and validate generated numerical data. Espana et al.

Espana et al. [4] used thermocouples to measure temperature response during Laser Engineered Net Shaping (LENS). Xiong et al. [5] and Ye et al. [6] used thermocouples and thermography in order to characterize temperature responses during Laser Engineering Net Shaping (LENS). Griffith et al. [7] have reported meaningful temperature responses as provided via thermocouples even after 42 layers of deposition during LENS. In these experiments, the thermocouple bead was exposed atop the substrate and the deposition occurred above and around it. Results of these studies indicate that, due to fast solidification process of the melt pool (which happens in milliseconds), it is hard to monitor its temperature by thermocouples alone.

Results of these studies indicate that, due to fast solidification process of the melt pool (which happens in milliseconds), it is hard to monitor its temperature by thermocouples alone. However, knowledge of thetemperature distribution outside the melt pool area allows one to better understand the microstructure evolution and to obtain more data for validating numerical simulations.Taylor [8] proposed a model to predict the effective thermal conductivity and absorptivity of a sintering surface and benchmarked the model with experimental data obtained via thermocouples exposed to either air or argon atmospheres. Shishkovsky [9] measured the temperature response around the melting area during the Selective Laser Sintering (SLS) of metal-polymer, intermetallic and ceramic materials using five thermocouples (temperature range 400-2800 ºC). In addition to these experimental approaches, there exist analytical and numerical methods for estimating the temperature response during SLM.

In addition to these experimental approaches, there exist analytical and numerical methods for estimating the temperature response during SLM. For instance, it is common to employ Rosenthal’s classical analytical solution to determine temperature distribution in media with a moving heat source problem [10]. Numerical methods, such as the Finite Element (FE) and Finite Difference (FD) methods, are used to solve more complex problems. Dong et al. [11] used the FE method to study the effects of phase transformation as well as to determine the influence of scanning speed, laser intensity and beam spot diameter on SLS parts during manufacture. Riedlbauer et al. [3] investigated the SLM melt pool temperature profile and its size by using a nonlinear FE method.There are different methods for varying local cooling rates during the SLM process, including the variation of process parameters such as laser power, laser speed, hatching space and thickness of each layer.

Variation of these parameters would impact the melt pool temperature and consequently alter cooling rates. However, it is possible that, upon changing process parameters during SLM, the porosity of the part can increase due to lack of fusion and/or melt pool instability. Hence, to ensure ideal deposition and solidification conditions, parameter sets, not individual parameters, should vary with time to ensure full-density and target microstructure. The combination of process parameters for ensuring near- or fully-dense parts depends on the utilized material, powder size and the building time.Since real-time process control of SLM is not fully available within the state-of-the-art, it is of interest to work with time-invariant process parameters during SLM. For constant process parameters, one can, for example, employ unique scan patterns to overcome effects of part size and geometry. Hence, in this study, effects of scan pattern on cooling rates during SLM will be studied.

Hence, in this study, effects of scan pattern on cooling rates during SLM will be studied. Also the effectiveness of changing them will be compared to changing the process parameters.AnalysisThis case study focused on the fabrication of 17-4 precipitation hardened (PH) stainless steel (SS). Numerical simulations were performed by first discretizing both space and time in order to solve the partial-differential, conservation equations for energy. The temperature through the solidified part and powder bed was assumed to be governed by a non-homogeneous heat equation with Fourier-type conduction.

The temperature through the solidified part and powder bed was assumed to be governed by a non-homogeneous heat equation with Fourier-type conduction. The heat equation governing temperature along a single layer can be written as:𝜌(𝑡)𝑐(𝑡)∂𝑇(𝑥, 𝑦, 𝑧, 𝑡) ∂t= 𝑘(𝑡)∇2𝑇(𝑥, 𝑦, 𝑧, 𝑡) + 𝑄(𝑥, 𝑦, 𝑡)(1)where 𝜌(𝑡) is temperature-dependent density of SS 17-4 PH, 𝑐(𝑡) is the temperature-dependent specific heat capacity of SS 17-4 PH, 𝑘(𝑡) is temperature-dependent thermal conductivity of SS 17-4 PH [12], T is temperature and Q is a volumetric heat source term assumed to be the irradiation flux density at the melt pool interface, i.e.:𝑄 = 𝛼𝐺 = 𝑓(𝑟)where α is the absorptivity (assumed constant) of the material and G is the laser power at a given location.

The laser power with Gaussian form is formulated as:𝐺 ≅2𝑃 𝜋𝑅2 exp (−2𝑟2 𝑅2 )where P is laser power, R denotes the effective laser beam radius at which the energy density reduces to 1 𝑒2⁄ at the center of the laser spot, r is the radial distance from the center of the laser spot [2]. The latent heat transfer associated with powder melting and melt pool solidification was modeled using:𝑞 = 𝑚𝐿where q is the energy released or absorbed during the change of phase, m is the mass of substance and L is the specific latent heat of fusion for a particular substance. The boundary conditions considered both convection and thermal radiation based on the exposed surface of the current layer to the surroundings, i.e.:𝑞c = ℎ(𝑇 − 𝑇0c) 𝑞r = 𝜎𝜀(𝑇4 − 𝑇0r4)where h is the heat transfer coefficient, σ is the Stefan-Boltzmann constant and ε is the emissivity. Also, 𝑇0c represents the temperature of gas inside the chamber and T0r represents the chamber surroundings temperature.

Also, 𝑇0c represents the temperature of gas inside the chamber and T0r represents the chamber surroundings temperature. The heat transfer coefficient was assumed as uniform and was calculated based on the velocity and the type of gas inside chamber. Given the small size of the melt pool relative to the SLM chamber, gray body/enclosure-type radiation behavior exists and Eq. (6) holds. The absorbed energy from the heat source it utilized for melting the powder or transfers through the powder bed, at a rate of 𝑞cond, or dissipates to the surroundings, i.e.:𝑄 = 𝑞 + 𝑞c + 𝑞r + 𝑞condBy solving Eq. 1, the temperature profile along a layer can be obtained in all points at all time. While adding a new layer, a previously-deposited layer has time to cool via powder-bed heat transfer or to the surrounding environment.Numerical SimulationNumerical simulations were performed using the commercially-available COMSOL 4.4 Multiphysics FE software.

The powder bed was assumed to be a homogenous and continuous medium and the laser heat source possesses a Gaussian intensity distribution (Eq. 3). The effective thermal conductivity of the powder bed was calculated using:𝑘pb = 𝑘s(1 − 𝜙)(2)(3)(4)(5)(6)(7)(7)where 𝑘pb and 𝑘s are the thermal conductivity of powder bed and solid powder, respectively. Porosity of powder bed, ϕ can be written as:𝜙 =𝜌s − 𝜌p 𝜌swhere 𝜌s and 𝜌s are densities of solid material and powder bed, respectively. For the current investigation, the porosity of the powder bed was 60%. A time step of Δt = 0.1 s was utilized for the current investigation, so it could capture all the transient responses. Free triangular elements were chosen for the meshing scheme. In order to better model the high heat flux transport near the laser/liquid interface, the element size in this region was selected as L ~ 20 μm.

In order to better model the high heat flux transport near the laser/liquid interface, the element size in this region was selected as L ~ 20 μm. Element sizes increased with distance from the high heat flux zone to a size of L = 100 μm, as shown in Fig. 1. The heat transfer coefficient was estimated based on the gas inside the chamber, its velocity and the geometry of the problem. Based on Reynolds number (3.04×104), Prandtl number (0.68) and Nusselt number (102), the heat transfer coefficient is 12.7 W/m²∙K. Finally, the initial condition for the powder and 17-4 stainless steel substrate was set as 20 ºC. Mesh independent results were obtained by reducing the size of the mesh to the size that results were not affected by the size of mesh. Parallel supercomputing resources were utilized to simulate the manufacturing process.Figure 1.

Parallel supercomputing resources were utilized to simulate the manufacturing process.Figure 1. Meshing scheme used for simulations showing one layer of powder and substrateExperimental Setup and ProcedureExperiments were conducted using a PHENIX PM 100 SLM system. The particular SLM system uses a fiber laser (Photonics) with Gaussian profile intensity, 50 W power limit, 1075 nm wavelength and spot size of 70 µm. All experiments were performed under argon atmosphere, pumped into the 200 L chamber at a rate of 167 cm3/s. Substrate temperature was monitored using, an embedded, pre-calibrated Type-C thermocouple (OMEGA, bead diameter ≅ 0.075 mm) as shown in Fig. 2. Temperature data were recorded using an ICP DAS I-7018 data logger(8)at a rate of 1 Hz. For the typical temperature range recorded, the maximum error associated with the thermocouples was 4.5 ˚C (±1%), while nominal error of the data logger was 0.2 %.The material used in the process was SS 17-4 PH powder.

The average particle size was . The 25 µm with a room-temperature thermal conductivity of approximately 12 W m ∙ K absorption coefficient, α, was assumed to be constant at 0.3 as suggested by [13, 14]. Two different sets of experiment were performed. The first experiment was performed without powder in the machine. The laser moved from the starting point, as shown in Fig. 2, to the edge of the plate a single time. During this experiment, three different laser speeds and powers were employed and these parameters are summarized in Table 1. The scan pattern was only performed once and then the substrate was allowed to cool back to room-temperature before a new set was performed.⁄Table 1. Process parameters utilized for nine laser scans along thermocouple-embedded substratePower\speed 25.0 W 37.5 W 50.0 W50 mm/s Exp. #1 Exp. #4 Exp. #7150 mm/s Exp. #2 Exp. #5 Exp. #8250 mm/s Exp. #3 Exp. #6 Exp. #9Figure 2.

#1 Exp. #4 Exp. #7150 mm/s Exp. #2 Exp. #5 Exp. #8250 mm/s Exp. #3 Exp. #6 Exp. #9Figure 2. Dimensions of thermocouple-embedded substrate with start point and laser path [15]The second experiment was designed for fabricating a thin wall structure atop the thermocouple-embedded substrate. The length and width of the wall was 50 mm and 2 mm, respectively. The thin wall was constructed using a laser scan pattern shown in Fig. 2. The process parameters were held constant with time and consisted of a 50 W laser power and 50 mm/s scan speed. All of the results from experiment were used to verify the numerical simulations. The substrate design and powder bed were ensured to match those related to the SLM machine/process utilized for validation experiments.Results and DiscussionThe temperature response at Point A (shown in Fig. 2), as measured experimentally and numerically for Exp. #7 (single laser pass) is shown in Fig. 3.

2), as measured experimentally and numerically for Exp. #7 (single laser pass) is shown in Fig. 3. For the numerical simulation of this experiment, the temperature under the substrate (Point A) was extracted after simulating heat transfer through powder and substrate. The results demonstrate that the temperature within the substrate during laser exposure does not surpass 60 ºC, and that the substrate cools within ~ 15 s. However, the experimental data appears to agree well with numerical data for all experiments conducted, i.e. Exps. #1 - #9. The numerical model approach is shown to be validated in Fig. 3.Figure 3. Variation of temperature at Point A with respect to time during Exp. 7 laser scanning.Figure 4 shows the simulated substrate temperature response during deposition of the SS 17-4 PH thin wall while using the same parameters from Exp. #7. The first layer of the thin wall is presented, and the multi-track scan pattern shown in Fig. 2 is utilized.

#7. The first layer of the thin wall is presented, and the multi-track scan pattern shown in Fig. 2 is utilized. The effect of scan pattern on the temperature response at Point A below the substrate is clearly evidenced by the multiple temperature peaks vs. time. Note that the experimental data is not oscillatory due to the employed temperature collection frequency; however, the general temperature rise agrees very well with the numerical simulation. During deposition of one layer, the temperature of the substrate has increased by 25 ºC, to a maximum temperature near 50 ºC. Note that the temperature cycling amplitude appears to reduce as the laser gets closer and closer to Point A.Figure 4. Variation of temperature at Point A with respect to time during SS 17-4 PH thin wall build using Exp. #7 process parameters and scan pattern.The effects of laser scan pattern on the temperature response was investigated by using each of the three scan patterns shown in Fig.

#7 process parameters and scan pattern.The effects of laser scan pattern on the temperature response was investigated by using each of the three scan patterns shown in Fig. 5 for the simulated SLM fabrication of a SS 17-4 PH thin wall. From Fig. 5, it may be seen that the laser can pass parallel to any edge of the wall, go back-and-forth or just in one direction parallel with the wall length.Figure 5. Scan patterns investigated: (a) alternating, (b) unidirectional/lengthwise and (c) unidirectional/widthwiseThe temperature response of Point A for all three scan patterns, using process parameters from Exp. #7, is shown in Fig. 6. From Fig. 6, it may be seen that each laser scan pattern provides for a unique temperature response at Point A within the substrate, and thus, a different temperature response in the surrounding powder bed and part. The results show that, like the trend observed in Fig.

The results show that, like the trend observed in Fig. 4, the temperature oscillations decrease in amplitude as the laser approaches the measurement point. Hence, the severity of temperature fluctuations is decreased as the bulk part temperature increases for a given layer. The alternating, zig-zag pattern (e.g. scan pattern #a) provided for the hottest peak temperatures and largest temperature amplitudes. Note that although the time invested per layer for all of three scan patterns is the same, the deposited part will experience completely different cooling and heating rates – as evidenced by the amplitude and frequency of the temperature oscillations at Point A. These different cooling rates will directly impact the encumbered part microstructure, and thus provides a means to control part thermal history without changing process parameters in real-time.Figure 6.

Variation of temperature at measurement location (Point A) with respect to time during SS 17-4 PH thin wall build using Exp. #7 process parameters and three different scan patternsThe heating and cooling rates (i.e. temperature time rate of change) while laser moving in one direction or moves back and forth presented in Fig. 7 and Fig. 8. The temperature was obtained through the simulation of thin wall build. Process parameters are same as Exp. #7. By comparing Fig. 7 and Fig. 8, it is obvious that temperature rate is strongly dependent on scan pattern. Scan pattern (b) nearly double the peak rate. By knowing the effects of scan pattern on thermal response, controlling the temperature rate become feasible.Figure 7. Temperature time rate of change at measurement location with respect to time using Exp. #7 process parameters with scan pattern (b)Figure 8. Temperature time rate of change at measurement location with respect to time using Exp.

#7 process parameters with scan pattern (b)Figure 8. Temperature time rate of change at measurement location with respect to time using Exp. #7 process parameters with scan pattern (a)The effects of geometry on temperature response and cooling rate were studied by fabricating a thin wall with various lengths via simulation. As shown in Fig. 9, a 2 mm thick wall with a length of 20 mm, 30 mm or 40 mm was constructed. The ‘alternating’ scan pattern, Fig. 5 a, and process parameters from Exp. #7 were utilized. The results for the first deposited layer are provided in Figs. 10-11.Figure 9. Thin wall dimensionsIn Figure 10, it is obvious that as the length of the thin wall decreases, for a given set of process parameters and scan pattern, the peak temperature increases. For length of sample reduce from 40 mm to 20 mm, the peak temperature increase from 50 ˚C to 57 ˚C.

For length of sample reduce from 40 mm to 20 mm, the peak temperature increase from 50 ˚C to 57 ˚C. the time of the build will decrease as the size decrease too, so we expect that heating rate would increase as length of part decrease.The heating/cooling rates are also influenced by the geometry of the part, as shown in Fig. 11. It may be seen that the cooling rate increases as the length of the wall decreases. However, the opposite is true for heating rate. The wall with length of 30 mm has the highest heating rate while heating rate of the other two walls is nearly the same.Figure 10. Variation of temperature at measurement location with respect to time using Exp. #7 process parameters with three different geometriesFigure 11. Temperature time rate of change at measurement location with respect to time using Exp.

#7 process parameters with three different geometriesFigure 11. Temperature time rate of change at measurement location with respect to time using Exp. #7 process parameters with three different geometriesConclusionsThe purpose of this research was to provide the tools to predict the heat transfer through the parts, powder and substrate. The feedback from model and thermocouples measurements could be used to decide whether the remaining portion of the powder is recyclable or not. Also results, temperature distribution, cooling and heating rate, are necessary for determining the microstructure formation. Finally effects of process parameters and built patterns on heat transfer were studied The model used to predict the temperature and rate of change of it over time for different geometry and scan pattern. It has shown that by changing these parameters, parts experience different thermal history.

It has shown that by changing these parameters, parts experience different thermal history. These results could be used to design the built pattern based on the geometry to get uniform response even with different geometry.It is understood from these simulations that by changing process parameters, scan pattern and geometry, completely different peak temperature can be obtained. Also each point in the part experiences different cooling rate. So the final product of SLM machine is non homogenous part which its mechanical properties is dependent on the process parameters, built pattern and the geometry. In order to solve the problem, the parts can be heat treated using classical methods. Also by designing unique built pattern based on geometry and adopting suitable process parameters more homogenous parts with same mechanical properties can be builtTo summarize, results could be divided into four groups, temperature distribution, heat transfer, cooling and heating rate.

Numerical modeling is capable to give information of entire thermal behavior. All un-melted powder remaining in the SLM machine is sought to be recycled for later use without detriment. However, un-melted powder near the part will undergo a degree of thermal cycling and may impact its recyclability.AcknowledgementsThis work was financially supported by Mississippi State University’s Office of Research and Economic Development (ORED) and Center for Advanced Vehicular Systems (CAVS). Authors acknowledge the resources provided by Mississippi State University’s High Performance Computing Collaboratory (HPC²).References[1] L. Bian, S. M. Thompson and N. Shamsaei, "Mechanical Properties and Microstructural Features of Direct Laser Deposited Ti-6Al-4V," JOM TMS, vol. 67, no. 3, pp. 629-638, 2014.[2] Y. Li and D. Gu, "Parametric analysis of thermal behavior during selective laser melting additive manufacturing of aluminum alloy powder," Materials and Design, vol. 63, pp.

Liu, "Characterisation of the Thermal Damage in a Martensitic Steel Substrate Consequent to Laser Cladding Process," Procedia Engineering, vol. 10, p. 2851–2856, 2011.[14] C. M. Taylor and T. H. C. Childs, "Thermal experiments in direct metal laser sintering," in Proceedings of 9 Assises Européennes du Prototypage Rapide and 10 European Conference on Rapid Prototyping and Manufacturing (EURO RP 2001), Paris, 2001.[15] M. Masoomi, E. Alaa, N. Shamsaei and S. Thompson, "Modeling, simulation and experimental validation of heat transfer during selective laser melting," in Proceedings of the ASME 2015 International Mechanical Engineering Congress & Exposition, Houston, 2015..

periments In Layered Electro-Photographic PrintingDenis Cormier, James Taylor, Kittinan Unnanon, Parikshit Kulkarni, and Harvey WestDepartment of Industrial EngineeringNorth Carolina State UniversityRaleigh, NC 7695AbstractElectro-photographic printing processes employed by products such as laser printers and photocopiers are commonly used to deposit and fuse thin layers of thermoplastic powder onto paper. This report describes preliminary experiments aimed at adapting the electro-photographic printing process for use as a layered manufacturing technique. 3-D electro-photographic printing holds considerable potential as an inexpensive freeform fabrication technique that is suitable for office environments. The possibilities for selective coloring are also discussed.IntroductionElectro-photographic printing processes employed by products such as laser printers and photocopiers are commonly used to print toner onto paper or plastic substrates.

Although the printed images are routinely referred to as "two-dimensional", they obviously do have a finite and measurable thickness. The toner used to print these images is a blend of several materials, the largest constituent of which is typically polystyrene. Given the fact that laser printed images have thickness, and that conventional toner is primarily a thermoplastic (polystyrene), the natural question arises as to whether the electro-photographic printing process can be adapted for use in a layered manufacturing technique. If such a process could make use of existing mass produced laser printer engines, then the process would potentially offer an extremely inexpensive approach to prototyping.

If such a process could make use of existing mass produced laser printer engines, then the process would potentially offer an extremely inexpensive approach to prototyping. Furthermore, the process would have considerable potential in the area of selective coloring.This paper reports on initial experiments designed to investigate the capabilities and potential of 3-D electro-photographic printing, and to identify the major technological hurdles that must be overcome in order for the process to become viable as a SFF technique. Issues pertaining to CAD model slicing for 3-D electro-photographic printing are also presented.Layered Electro-Photographic PrintingOne of the earliest references to the concept of layered manufacturing via electro- photographic printing can be found in Bynum's patent (Bynum, 1992). The patent describes four unrelated layered manufacturing techniques, one of which relates to electro-photographic printing.

The patent describes four unrelated layered manufacturing techniques, one of which relates to electro-photographic printing. Bynum’s concept involves using an electro-photographic process to selectively deposit a thermoplastic toner onto a continuous teflon coated belt. The toner on the belt is heated to the point that it becomes “tacky”. The belt is then indexed such that the printed image is registered267above the previously deposited layers. Finally, a platen presses the tacky printed image onto the top of the previously printed images. It does not appear that a functioning apparatus was ever built, or that experimental details have ever been published.Johnson (1994) describes an interesting process that he terms Particle Deposition Fabrication. This process is similar to Bynum’s concept in the sense that thermoplastic toner powder is selectively printed and fused. However, the method of depositing the toner is slightly different.

This process is similar to Bynum’s concept in the sense that thermoplastic toner powder is selectively printed and fused. However, the method of depositing the toner is slightly different. With this process, a magnetic core particle carrier drum has a thin uniform coating of magnetic toner particles applied to it. Beneath the carrier drum lies a matrix of control electrode apertures. Each aperture is individually addressable such that electromagnetic fields that induce toner to jump off of the carrier drum and onto the printed object can be selectively turned on and off. This technique is used to selectively deposit toner particles that are subsequently thermally fused to the substrate. As is the case with Bynum's approach, it does not appear that a functioning apparatus using this approach has yet been built.Outside of the two conceptual techniques described by Bynum and Johnson, virtually no details on layered electrophotographic printing can be found in the literature.

Despite the lack of published data, if a process could be developed that employed conventional electrophotographic print engines, then the process would hold considerable promise as an inexpensive SFF technique. Much of the promise would derive from the fact that the most expensive component (i.e. the print engine) is already mass produced and therefore is quite inexpensive. Furthermore, color laser printers are becoming quite inexpensive. For example, the Hewlett Packard LaserJet 4500 color laser printer currently has a street price of approximately $2,500 (U.S.).

Furthermore, color laser printers are becoming quite inexpensive. For example, the Hewlett Packard LaserJet 4500 color laser printer currently has a street price of approximately $2,500 (U.S.). This indicates that a color layered electrophotographic printing process would also be relatively inexpensive.Due to the fact that evidence of an actual functioning apparatus could not be located, a series of preliminary experiments designed to shed light on the challenges associated with developing a viable layered electrophotographic printing process were conducted.Initial Printing ExperimentsThe initial experiment used a Hewlett Packard Laserjet 2100 laser printer. First, a layer of polyvinyl alcohol (PVOH) film was affixed to a paper substrate. The PVOH sheet was then fed through the laser printer 30 times while the same image of a simple rectangle was printed onto the same location of the sheet. Upon completion, the paper was placed in warm water.

Upon completion, the paper was placed in warm water. Since PVOH is soluble in water, the PVOH beneath the printed image dissolved, thus allowing the 30 layer thick printed rectangle to release from the paper. This basic process was repeated numerous times.Although the initial experiments worked well, the resulting parts were quite brittle. According to Pickett (2000), toners come in both magnetic and non-magnetic formulations. The choice depends on which type of print engine is used. The HP 2100 laser printer uses the more common magnetic toner formulation which consists of a polystyrene-acrylic copolymer (approximately 60% by weight) and ferrous oxide (approximately 40% by weight) blend (Hewlett Packard, 2000). The polystyrene-acrylic copolymer has a glass transition temperature (Tg) that is approximately 55˚C. Since this Tg value is above room temperature (i.e.

The polystyrene-acrylic copolymer has a glass transition temperature (Tg) that is approximately 55˚C. Since this Tg value is above room temperature (i.e. it is268relatively brittle at room temperature), the powder can be inexpensively mass produced via conventional ball milling. Unfortunately, the polystyrene acrylic copolymer toner that is specially formulated to facilitate mass production also leads to brittle laser printed objects. Thus, the formulation of toners with more desirable mechanical characteristics is an open research issue for this process.Any efforts to formulate new toners must obviously be compatible with the process that is used to print the toner. Pickett (2000) indicates that the rollers used to pick up toner within a toner cartridge have a texture that is specifically designed to pick up a single layer of toner particles. It is desirable, therefore, for newly formulated toners to have similar size and shape characteristics as conventional toner.

It is desirable, therefore, for newly formulated toners to have similar size and shape characteristics as conventional toner. In order to examine the basic size and shape of conventional toner, a sample was placed beneath a microscope (Figure 1-a). The scale in this figure is 10 _m per division, thus toner particles appear to be spherical with a diameter of 10 – 15 _m. For comparison, a sample of high density polyethylene (HDPE) was placed under the microscope (Figure 1-b). Although some of the particles tended to clump together, the individual particles are approximately 10 _m in diameter and also have a predominantly spherical shape. Based upon this fact, a toner cartridge was emptied and cleaned out with the assistance of a local toner cartridge remanufacturing firm. The conventional toner was replaced with the HDPE powder, and the printing experiment was repeated.

The conventional toner was replaced with the HDPE powder, and the printing experiment was repeated. Although the fusion temperature was not adjusted lower for the HDPE, the laser printer printed this powder without any difficulty.Selective Coloring ExperimentsBased upon the initial demonstration of feasibility, a second set of experiments involving selective coloring was attempted. This experiment used a Hewlett Packard LaserJet 4500 color laser printer. As was the case with the initial trials, PVOH film was attached to a sheet of paper that was subsequently passed through the color laser printer 30 times while the same image was printed. Initially, solid rectangles in a variety of colors were printed and released from the PVOH substrate.

Initially, solid rectangles in a variety of colors were printed and released from the PVOH substrate. In subsequent experiments, multi-color images such as yellow and black bulls-eyes were printed (Figure 2).(a)(b)Figure 1 - (a) Toner Particles; and (b) HDPE Particles (10 microns per division)269Figure 2 - Yellow and Black Bulls-EyeInitial trials with color printing revealed an interesting phenomenon. With each successive layer that was deposited, the parts grew darker and darker. It was hypothesized that the color toners are formulated such that the color and luminescence of the printed image account for the backlighting that occurs when light bounces off of the white paper after having passed through the toner. In order to test this hypothesis, two experiments were performed. In the first experiment, a series of red rectangles were printed, each of which had more layers than the previous rectangle.

In order to test this hypothesis, two experiments were performed. In the first experiment, a series of red rectangles were printed, each of which had more layers than the previous rectangle. Images of these samples were captured using a flatbed scanner, and the color and luminescence were measured using Adobe PhotoShop. The experiment revealed that the color was independent of the number of layers, but that luminescence decreased as the number of layers increased.For the second experiment, a sample of experimental white toner was acquired from a toner supplier. The printer's black toner was replaced with the white toner, and a set of rectangles with varying numbers of layers were again printed. With this set, however, all but the top layer were printed using white toner. Only the final (top) layer was printed in red. Again, the samples were scanned and analyzed using PhotoShop.

With this set, however, all but the top layer were printed using white toner. Only the final (top) layer was printed in red. Again, the samples were scanned and analyzed using PhotoShop. It was found that both color and luminescence remained constant regardless of the number of layers when the samples consisted of a white "core" on top of which a color "skin" was printed. More complete details on these experiments can be found in Cormier et al. (2000).Build Speed and ResolutionOne of the initial concerns with layered electrophotographic printing was build speed. Intuition suggests that the printed layers are so thin that the process would be too slow to be of practical use. In order to estimate the build speed of the process, a laser profilometer was used to measure the thickness of 20 laser printed rectangles. The average layer thickness was 8.35 _m with a standard deviation of 1.67 _m.

The average layer thickness was 8.35 _m with a standard deviation of 1.67 _m. Mid-level laser printers that currently sell for approximately $700 (US) have a typical print speed of 14 pages per minute. Note that this rate is based upon printed pages that are 8_˝ wide x 11˝ long. If 8.35 _m per layer of material is deposited at a rate of 14 layers per minute, then a layered electrophotographic printer built using a mid-level print engine would build parts at a rate of approximately 7.014 mm/hr (0.276 in/hr)..000835mm hr«14layers min60min hr=.7014mm hr270Note that this build rate is based upon parts that require the full 11˝ of travel per layer. As it turns out, this deposition rate is considerably faster than intuition would suggest, and it is well within the range of existing SFF processes for a part of this size.

As it turns out, this deposition rate is considerably faster than intuition would suggest, and it is well within the range of existing SFF processes for a part of this size. Smaller parts would not require the print carriage to travel as far, thus the build speed would be correspondingly faster for these parts.With regards to resolution, low-end laser printer engines currently have resolutions of 600 dots per inch, while mid-level laser printer engines have resolutions of 1200 dots per inch. These two resolutions equate to 0.0017 inches per dot (0.042 mm/dot) and 0.0008 inches per dot (0.021 mm/dot) respectively. If perfect registration from one layer to the next is achieved, then these resolution figures represent the theoretical minimum feature size that can be achieved.

If perfect registration from one layer to the next is achieved, then these resolution figures represent the theoretical minimum feature size that can be achieved. As is the case with most commercial SFF processes, however, the actual achievable feature size is more a function of the material properties than the process capabilities.Slicing for Layered Electro-photographic PrintingDue to the fact that layered electro-photographic printing uses a substantially different method to deposit the material, a slightly different approach to slicing is necessary. Most SFF techniques involve motion control of a laser, printhead, or some other device that produces solid layers in a raster fashion. In the case of layered electro-photographic printing, an image is printed, thus the slicing software must produce a graphic image of the slice to be printed.

In the case of layered electro-photographic printing, an image is printed, thus the slicing software must produce a graphic image of the slice to be printed. Furthermore, images of both the build and support material must be constructed.In order to produce the required images, two different slicing algorithms have been developed and implemented. The first slices STL files via conventional slicing techniques. The second employs a direct slicing approach using the SolidWorks CAD system. Both implementations output a text file containing slice contour data for both build and support regions within each layer. An OpenGL program has been written in Visual C++ that reads in the slice contour data and graphically renders the build and support material images to the appropriate scale. At the present time, these images can be manually directed to a printer one at a time.

At the present time, these images can be manually directed to a printer one at a time. The next step will be to select an archival file format for storing the slice image data.The direct CAD slicing algorithm employed within SolidWorks uses the followingapproach:Procedure DirectSlice For each sliceCreate a slice cube at the correct Z-height Intersect the slice cube with the part For each body in the intersection volumeGet the number of loops in the current intersection volume For each loop in the current intersection volume Get the loop type (i.e. inner or outer loop) For each segment in the loop If the segment is a line Record the line end pointsElse271Tesselate the loop Record the tessellation points in the correct orderNext SegmentNext LoopNext BodyNext Slice End Procedure DirectSliceThis procedure has been implemented using the SolidWorks API.

Note that a check must be made to ensure that outer loops are oriented counter-clockwise, whereas inner loops are oriented clockwise. This convention is necessary so that the images can be properly constructed. Figure 3(a) shows a screen dump from the SolidWorks window showing the slice outline for a connecting rod that is in the process of being sliced.Once the slice contour data file is generated, the images can be generated. A Visual C++ program that uses OpenGL graphics routines has been written that reads in the slice contour data and renders it to the computer screen. In order to do this, it is important to render outer and inner loops in the correct sequence so that inner loops are properly rendered.

In order to do this, it is important to render outer and inner loops in the correct sequence so that inner loops are properly rendered. Figure 3(b) shows the slice contour image that has been generated for the connecting rod example from Figure 3(a).(a)(b)Figure 3 – (a) Direct Slicing Screen Dump; (b) Rendered Slice ImageOpen Research IssuesBased upon the initial preliminary experimentation, a set of open research issues has been compiled. In order to address these issues, it is first necessary to understand the basic operating principles of a typical electro-photographic print engine, such as the one shown in Figure 4.272LaserToner CartridgeOPC DrumFuserPaperTransfer RollerFigure 4 - Schematic of Typical Laser Printer EngineThe OPC (organic photo-conductive) drum is a cylinder with a special coating that is capable of locally holding a static charge. A laser or LED source is used to "write" a latent image onto the OPC drum as it rotates.

A laser or LED source is used to "write" a latent image onto the OPC drum as it rotates. This image is initially nothing more than selectively applied electrostatic charge. Within the toner cartridge, a thin coating of toner powder is applied to another rotating drum. As the toner cartridge drum and OPC drum rotate near one another, toner is induced to "jump" onto the OPC drum by the selectively applied static charge on the surface of the OPC drum. The OPC drum continues to rotate until the toner comes in contact with the paper. A transfer roller beneath the sheet of paper induces an even stronger force on the toner particles, thus inducing them to become deposited on the paper. As the paper exits the machine, the toner is thermally fused to the paper.With regards to layered electro-photographic printing, perhaps the most significant technological challenge lies in inducing the printed image to leave the OPC drum and to be deposited onto the build platform.

With a conventional laser printer, the paper sheet is thin enough that a transfer roller induces sufficient force through the paper. With layered manufacturing, the induced force will decrease as the part thickness increases. Therefore, an alternative method of transferring the toner powder must be employed. As stated previously, Bynum (1992) proposes heating the toner to the point that it becomes "tacky" and then mechanically pressing it onto the workpiece via a platen. Johnson (1992) identifies several other transfer techniques that have been used with printing processes, and which may have applicability for layered electro-photographic printing. These techniques include magnetization and triboelectrification.273Another issue that has not yet been mentioned is that of the support material. The approach that is being considered for the experimental apparatus is to use a second print engine that prints PVOH powder that is simply dissolved in water upon completion of printing.

The approach that is being considered for the experimental apparatus is to use a second print engine that prints PVOH powder that is simply dissolved in water upon completion of printing. The challenge will be to formulate a PVOH material that has the same particle size as ordinary toner, and which has similar fusion properties as the toner. Furthermore, adhesion tests between polystyrene toner and PVOH will need to be run.SummaryThis paper has reported on preliminary tests designed to determine the basic feasibility of layered electro-photographic printing. Initial test results are encouraging, however, the investigation has also revealed several substantial technological challenges.Efforts are currently underway to build a monochrome layered electro-photographic test apparatus. Every attempt is being made to design the initial prototype so that components and sub-systems can be interchanged while different configurations are evaluated.

Every attempt is being made to design the initial prototype so that components and sub-systems can be interchanged while different configurations are evaluated. Collaborative efforts with Materials Science faculty are also underway to investigate issues pertaining to alternative toner formulation and their manufacture.ReferencesBynum, D. (1992) "Automated manufacturing system using thin sections," United States Patent # 5,088,047.Cormier, D., Taylor, J., Kulkarni, P., and West, H. (2000) "An Investigation of Selective Coloring With 3-D Laser Printing" submitted to the Rapid Prototyping Journal.Hewlett Packard (2000) http://www.hp.com/cposupport/printers/support_doc/bp101534.htmlJohnson, J (1994) Principles of Computer Automated Fabrication, Palatino Press, Irvine CA.Johnson, J (1992) Principles of Nonimpact Printing, Palatino Press, Irvine, CA.Pickett, J. (2000) Static Controls Inc. personal communication.274.

5. 6. Dept. of Computer Science, University of Notre Dame 7. Dept. of Chemical and Biomolecular Engineering, University of Notre Dame Johnson & Johnson Co. ITAMCOAbstract: The additive manufacturing (AM) process Direct Metal Laser Sintering (DMLS) uses a layer-by- layer workflow to build complex architecture metal structures in low-volumes. The primary process failure mechanism is a thermal stress driven thermal distortion that yields out-of-tolerance manufacture or complete process failure. However, DMLS design experts have developed heuristic rules to optimize the part orientation and support structure to reduce the likelihood of failure. We believe that experts innately attempt to minimize the design metrics of support volume (V), support-to-part surface area (A), maximal cross-sectional area of the slicing planes (X), parallelism of part faces with the recoater blade (P), and part height (H); however, it is unclear what relative weighting of each metric the expert uses.

This manuscript details an interactive expert survey, the statistical analysis of the survey responses, and the synthesis of an automatic algorithm for part orientation based on survey data. We received responses from 18 experts and 151 total part orientation responses. The median survey respondent had greater than four years of DMLS experience. Our analysis shows that the expert attempts to minimize metric V the most, metric X the second most, and metric H the third most; experts put essentially no weight on metrics A and P. The manuscript concludes with two orientation design studies where the expert survey responses are used in a least squares minimization algorithm to automatically orient the part for DMLS manufacture. As a comparison set, novice users were instructed to orient the parts for best DMLS printing success without using the tool and required multiple attempts to successfully print the test parts.

As a comparison set, novice users were instructed to orient the parts for best DMLS printing success without using the tool and required multiple attempts to successfully print the test parts. The automatically oriented parts failed on our first iteration of the code. The manuscript concludes with our proposed modifications to the code to improve results.Keywords: additive manufacturing, DMLS, optimization, build orientation, support structure, heuristic learning, expert survey1. IntroductionDirect metal laser sintering (DMLS) has been one of the primary success stories of the adoption of Additive Manufacturing (AM) tools into production cycles.

IntroductionDirect metal laser sintering (DMLS) has been one of the primary success stories of the adoption of Additive Manufacturing (AM) tools into production cycles. DMLS uses a layer-by-layer build cycle in which at each layer (Figure 1): 1) a recoater blade spreads a thin layer (20 – 50 µm) of1525metal powder at the melt plane; 2) radiant energy is applied to the melt plane by a raster laser selectively melting a two-dimensional (2-D) pattern; and 3) the part is indexed downward to accommodate a new layer of material. In this manner, complex three-dimensional (3-D) parts can be made for relatively low cost in a low-volume production setting (1 – 1,000 units) [1]. This manufacturing modality has been used in a variety of high performance aerospace industries including GE aircraft engines [2] and the SpaceX rocket engine chamber.

This manufacturing modality has been used in a variety of high performance aerospace industries including GE aircraft engines [2] and the SpaceX rocket engine chamber. Analogously, the orthopedics industry requires high-performance materials with different performance metrics (material purity, biocompatibility, and wear resistance are paramount), and are carefully considering the transition from traditional manufacturing processes (investment casting) to AM to reduce the per part manufacturing cost associated with the labor intensity, materials validation, and patient specific needs of their manufacturing processes. Further, market pressures are driving orthopedic manufacturers to greater reliance on mass customization and batch production, where the advantages of AM over other processes are greatest.Figure 1.

Further, market pressures are driving orthopedic manufacturers to greater reliance on mass customization and batch production, where the advantages of AM over other processes are greatest.Figure 1. Schematic of the DMLS printing process.AM manufacturing engineers utilize material supports to mechanically support overhanging features, but also, and perhaps more importantly, create sintered preferred conduction heat transfer paths to more efficiently purge heat from the recently sintered surface, so that cooling is more uniform across the build platform. Although necessary, any support material is inefficient as it requires longer build times, more material waste, and the need to remove the supports via fracture or machining in a post-process step [3]. Furthermore, support design is not a simple, binary (yes or no) decision as there are many competing design objectives that confound the rational decision on a ‘good’ design.

Furthermore, support design is not a simple, binary (yes or no) decision as there are many competing design objectives that confound the rational decision on a ‘good’ design. For example, part orientation is a critical parameter that influences the amount of support needed [4]. Support structures span a range of design features [3], and some overhanging features have an aspect ratio such that a support structure has a non-appreciable benefit.Clearly there is an economic design tradeoff between an excess of support material for efficient1526heat transfer and high-fidelity manufacturing and minimal support to reduce manufacturing costs. However, the design rules for support design are largely ad hoc and often require trial and error iteration, increasing development costs and materials and energy waste.

However, the design rules for support design are largely ad hoc and often require trial and error iteration, increasing development costs and materials and energy waste. Expert DMLS manufacturing engineers have an intuitive knowledge of the appropriate part orientation and shape, structure, and location of supports for heat spreading, but this innate knowledge is not easily disseminable at the rate at which DMLS, and AM in general, is expanding in market share. Furthermore, good design principles are not efficiently transferred to the novice and requires an experiential knowledge: experience is expensive as DMLS instruments cost over 500,000 USD and quality metallic powders cost over 120 USD per kilogram.

Existing support design software packages such as Materialise Magics SG+™ (Magics™) have shown promise to automatically design support structures and will even optimize part orientation based on a user-selected levels of z-height, xy-axis projection, support surface, and maximum slice area. However, it is unclear what the correct levels should be.In this manuscript, a method is detailed to codify the workpiece orientation problem in DMLS manufacturing design. Other design considerations, such as support design or operating parameter selection, although critically important, are outside the scope of this work. Theorientation design space can be defined by three Euler anglesv    :    180 ,180,    180 ,180 , and    180 ,180 . There are an infinite number of possibleorientations; from a manufacturing perspective, each orientation will have an associated quality (e.g.

There are an infinite number of possibleorientations; from a manufacturing perspective, each orientation will have an associated quality (e.g. thermal distortion or surface roughness), requisite designed countermeasures (e.g. support locations and reinforcing members), and manufacturing cost. From discussions with expert DMLS practitioners and a survey of the literature, it was determined that proper orientation design is a weighted minimization of the following metrics:Support Volume,( )V v : Supports are necessary to improve quality; however, excess supportvolume wastes material and machine time.Part-to-Support Surface Area, ( )Part-to-Support Surface Area, ( )have a more efficient thermal path to purge melt pool heat; however, increasing the surface area increases the support removal cost.Maximum Slice Cross-Sectional Area,( )X v : Melting induced shrinkage creates a contractilestrain in a given layer.

The larger the cross-sectional area, the larger the gross deformation;hence, there are advantages to designing orientations such that( )X v is small.Parallelism with Recoater,( )P v : Experts have defined a heuristic that major faces of the partshould not be within 10○ of parallel with the recoater blade as the blade incrementally deforms the part during each recoat step if the part-blade contact interface is large.1527 Part Height,( )H v : An orientation that results in a larger part height requires more layers,hence longer build times and more powder use.However, depending on the part, the different metrics can be antagonistic; for example,decreasing( )V v may increase( )H v . Thus, the definition of the respective weighting vector,w w w w w w , 2 5,,,134, of a single weighted objective function,Jxw(1)wherex V S X P H,,,,T, is not obvious.

In this work, a survey of DMLS experts andcorresponding method to interpret survey responses is described. Once an estimate of the weighting vector, ˆw , is found, the optimization algorithm in Eqn. 1.2 can be applied.ˆ min J vˆ xw(2)where ˆˆ J Wxis an estimated cost. This manuscript describes an expert survey to find and thenapply the appropriate weightings, ˆw . The general method uses the workflow diagrammed in Figure 2, the main methods will be detailed in Section 2. The survey results and application to two different test artifacts is given in Section 3. The paper concludes with a discussion on why the first iteration of the optimization based on the survey data failed and how this algorithm could be improved in the future (Section 4).Figure 2. Overview of the heuristic based orientation optimization.15282.

Overview of the heuristic based orientation optimization.15282. MethodsThe method for defining a heuristics-based orientation optimization algorithm is divided into two parts (Figure 2): the definition of a heuristic by survey administration and processing and the application of the optimization algorithm using the defined heuristic.2.1.PreliminariesIn general, the digital file that represents a given part is given by a .stl file, which represents the part as a surface composed of triangular facets. In simplest form, the .stl file is a list of facet parameters; each facet is parameterized by the unit vector normal to the facet and the facet vertices(Figure 3a).

In simplest form, the .stl file is a list of facet parameters; each facet is parameterized by the unit vector normal to the facet and the facet vertices(Figure 3a). For an N facet .stl file, there are two important matrices,V 33N and3N N  ,which are composed as follows:V      v i,1 xv i,2xv i,3xv i,1v i,2v i,3yyyv i,1 zv i,2zv i,3z      ;N    n i xn iyn iz   .(3)A given .stl file is rotated via a mathematical mapping operation on the matrices V and N . Herethe convention is used of rotating around intrinsic axes in the sequencex y z , although other ,',''equivalent rotation conventions exist (Fig.

Herethe convention is used of rotating around intrinsic axes in the sequencex y z , although other ,',''equivalent rotation conventions exist (Fig. Figure 3b): matrices V and N are post-multiplied bythe matrixR   ,,R z  Ry R   xwhereR xRyR z( ) ( ) ( ) 010           cos                  sin0 cos cos0 sin0 sin 0 1cos 0sin0 cos0sin cos    sin0001(4)1529Figure 3. .stl file conventions and manipulations. (a) Standard triangular facet partrepresentation by an .stl file. (b) Intrinsicxy '  z''angle convention for solid body rotation.Conventionally, a facet has support material beneath it is determined by the angle between the normal vectors of the facet and a unit vector normal to the build plane, {0,0,1}.2.2.

Definition of HeuristicSurvey Design To best understand how expert DMLS designers orient a workpiece, an interactive survey was administered. Input was solicited from July 15 2016 through January 3 2017, from individuals with expertise in AM, in general, by contacting members of the America Makes membership, colleagues with AM expertise, and using the professional surveying service Maven. The survey participants provided answers via a web portal written in Java that was divided into two parts: participant information collection and interactive part orientation.Participant information collection. Participants provided key information on their background and expertise through a set of questions (Table 1).

Participants provided key information on their background and expertise through a set of questions (Table 1). The questions are designed to help us understand the distribution in expertise level of the participant and what are the differences in orientation design between an expert DMLS engineer and a novice DMLS engineer, defined by greater than three years and less than three years of experience, respectively.Table 1.

Inconel), F) Refractory alloys (e.g.CoCr alloys), G) Other (open free answerbox if checked)Part orientation survey. The part orientation interface of the survey was designed to resemble a standard computer aided design (CAD) software package. The user can see a 3D rendering of1531the part, a visualization of the support structure, the build plane, and the recoater blade direction (Figure 4). Users are prompted to "orient the part for best manufacturability by the DMLS process," and provided details on the interface; additionally, users are asked to consider that the material is the material specified in Q8 if they were able to answer Q8 and aluminum if they were unable to answer Q8. The support structure is automatically generated using a standard heuristic that a facet should be supported if the angle of the facet is < 35○ and if the facet lies in the bottom 5% of the part.Figure 4.

Screen shot of the interactive survey demonstrating a case part (gray), support structure (red lines), build plate size (white plane), and recoater blade direction (blue triangle).The survey has a catalog of 20 parts, which were compiled from sample parts from the aerospace, automotive, orthopedic industries, and a few miscellaneous parts. In composite, the set includes a variety of features: both flat and contoured surfaces, large and small parts, and small and large holes.Data Collection and Extraction of Metrics. All data is logged by the survey software. Each survey respondent has provided an orientationset v , each will have a corresponding set of metrics (e.g.( )V v ) which are calculated by usingthe .stl. processor detailed in our previous manuscripts [5]. Each metric then has to be normalized to account for the large difference in part size and form factor of all possible parts.

processor detailed in our previous manuscripts [5]. Each metric then has to be normalized to account for the large difference in part size and form factor of all possible parts. The normalization is computed by the following1532ˆ( ) V vˆ( ) A vˆ ( ) X vˆ( ) P vˆ ( ) H vV V( ) vV  V minmin A Amax ( ) vA A minmax X ( ) v Xmin X X minminmax P ( ) v P max H ( ) v HP min P min H H  minmaxmin(5)where each minimum and maximum metric is computed via an optimization algorithm that searches for the maximum and minimum metrics by searching through the set of all possible angles v and computes the metrics at each angle. The optimization uses the dividing rectangles (DIRECT) algorithm to sample the design space and then converge to the optimal solution, which is defined as the minimal or maximal metric.

The optimization uses the dividing rectangles (DIRECT) algorithm to sample the design space and then converge to the optimal solution, which is defined as the minimal or maximal metric. Thus, the DIRECT method has to be run ten times for each part in the library of survey parts to generate all ten maximums and minimums of the metrics.Cost Function Design. Ultimately, all this data should be used to get a sense of what a DMLS designer weights when designing an orientation. The goal was to determine the objective function that is in the head of the designer.

The goal was to determine the objective function that is in the head of the designer. To do this, a constrained least squares estimator was used to extract objective function metrics from the data.ˆ ˆ Xwmin ˆ w2 subject to(6)ˆ w11 andˆ w0WhereˆX is the matrix of orientation data from the N responses organized by metric typeˆ Xˆ  V 1  ˆ V  2      ˆ V  Nˆ A 1 ˆ A 2  ˆ A Nˆ X ˆ X  ˆ X12Nˆ ˆ P H 1 ˆ ˆ P H 2   ˆ ˆ P H N12N      .(7)ˆw is a column vector of the unknown weightings on each metric type. In words, Eqn. (6) states1533“find the set of unknown weights ˆw that minimizes the product of the data and the weights, but subject two constraints: the weights must be positive and they must sum up to 1.” The requirement that the weights must sum up to 1 prevents the answer from being ˆ 0w  . This constrained least squares estimator was implemented using the Matlab function fmincon.2.3.

This constrained least squares estimator was implemented using the Matlab function fmincon.2.3. Heuristic Based Optimization.Optimization. ˆw can now be used to optimally orient any new part. First, a part is imported. Second,minimal and maximal values are found for metricsˆ V A X P H . Third, the objective functionˆˆˆˆ,,,,is minimized.ˆ x min ( , ,  ˆ w, )  ,(8)where ˆ(x    is a vector for the normalized metrics,,)ˆ V A X P H for Euler anglesˆˆˆˆ,,,,,  . This yields angles,,   for printing. This algorithm was applied to two parts: the,rectangular bar and the handle. Optimization of (8) was done by making use of the Direct Method heuristic optimization algorithm. DIRECT optimization algorithm is one type of derivative-free algorithm first introduced in [6], which is capable of solving difficult global optimization problems with bound constraints and real-valued objective function.

The history and development of DIRECT method is discussed in [7]. The name DIRECT is short for “DIviding RECTangles” because the DIRECT algorithm keeps dividing the current domain into new hyper- rectangles and identifying potentially optimal hyper-rectangles from the new hyper-rectangles. Figure 5 shows the first three iterations of the DIRECT algorithm [8]. At the beginning of the first iteration, or initialization of the algorithm the search domain is divided along all dimensions. Note that the division starts with the dimension with the best function value so that the best function value is left in the largest space. Then the potentially optimal hyper-rectangles are found by applying the lemmas developed from the Lipschitzian optimization [8]. Next, the potentially optimal hyper-rectangles are divided along the longest dimension.

Next, the potentially optimal hyper-rectangles are divided along the longest dimension. If the lengths along all dimension are the same, the optimal hyper-rectangles are divided along all dimensions, starting with the dimension with the best function value, as what was done in the initialization. The optimization algorithm keeps finding new potentially optimal hyper-rectangles and dividing those hyper-rectangles until the maximum number of divisions is reached.1534Figure 5. The schematic of the initial three iterations of the DIRECT optimization algorithm (Image has been copied from [8]).Case Study Evaluation A Case Study was done to validate the effectiveness of the optimization based on an expert survey. The study was comprised of two parts, a 10 x 10 x 150 mm rectangular bar (Figure 6a) and a contoured handle (Figure 6b) with a bounding box of 59 x 16 x 162 mm.

The study was comprised of two parts, a 10 x 10 x 150 mm rectangular bar (Figure 6a) and a contoured handle (Figure 6b) with a bounding box of 59 x 16 x 162 mm. The bar was selected because it was a simple structure with linear features, while the contoured handle represented a complex structure with contoured features.The intent of the case study was to compare the build setup as performed by two novice operators (a baseline condition) to that of the expert survey based orientation optimization. Parts were printed with 316L at 20 µm layer thickness on an EOS M290™ machine using the EOS provided parameter set. The success metrics are 1) Number of Attempts and 2) Maximum Distortion. Distortion measurements were made with a coordinate measuring machine (CMM) both on- and off- the build plate. Removal of parts from the build plate was done through wire electrical discharge machining.1535(a) (b)Figure 6. Case Study parts.

Removal of parts from the build plate was done through wire electrical discharge machining.1535(a) (b)Figure 6. Case Study parts. a) rectangular bar; b) contoured handle.Baseline ConditionsIn the baseline condition, two novice operators were used at Johnson & Johnson for the build set up utilizing their limited knowledge of the process. Operator 1 was a technician with one year of machine operation experience, but limited build preparation experience. Operator 2 was a high school intern with a month of polymer additive manufacturing experience and no metal additive manufacturing experience. Each operator was given basic instruction on Magics™, similar to that provided by EOS in their machine basic training curriculum.Each operator was given the .stl files with parts orientated as depicted in Figure 6 and the default Johnson & Johnson Magics™ support structure parameters.

Each individual was instructed to orientate the part and modify support structure as needed for successful printing, as defined by no part-support delamination through visual inspection. The operators were blind to the other’s attempts.3. Results3.1. Expert Survey-based optimization resultsThere were 18 respondents and a total of 151 different part orientations completed for an average of 8.4 orientations per respondent. The respondents spanned industry, national labs and agencies, and academia. The histograms in Figure 7 demonstrate the experience level corresponding to each orientation response.1536Figure 7. Experience level of the respondents. The frequency is the number of orientations for each experience level. As some respondents provided more orientations than others, I decided to present the experience tied to each orientation.Figure 8 is a boxplot of the metrics for all responses.

As some respondents provided more orientations than others, I decided to present the experience tied to each orientation.Figure 8 is a boxplot of the metrics for all responses. In general, respondents ensure that their orientations led to smaller support volumes (V), maximal cross-sectional area of a plane (X), and parallelism of faces with recoater blade. Respondents did not minimize the support-to-part surface area (A) and part height (H).Figure 8. Box plot of the orientation design metrics of all responses (N = 151).1537There are some reasonably interesting trends observed by breaking the orientation design metrics down by experience level (Figure 9). However, one should be careful to draw too much insight from the trends as the main effects of the different experience categories do not have a statistically significant effect on the metrics (all p > 0.05).

In general, the more experiences DMLS practitioner allows a larger support volume and height, but penalizes the maximal cross sectional area of a plane. Trends are even less noticeable for level of DMLS duty cycle in one’s job (Figure 10) or primary material used (Figure 11).Figure 9. Box plot of metrics separated by DMLS experience level. All differences arenot statistically significant.1538Figure 10. Box plot of metrics separated by DMLS Duty Level. All differences are notstatistically significant.Figure 11. Box plot of metrics separated by DMLS Material used. All differences arenot statistically significant.This constrained least squares estimator from Eqn. (6) was implemented using the matlabfunction fmincon.

All differences arenot statistically significant.This constrained least squares estimator from Eqn. (6) was implemented using the matlabfunction fmincon. The result is the following weights on V, A, X, P, H, respectively.1539ˆ w 0.5232 0 0.3571 0 0.1197 T (9)What this result ultimately says is that the respondents weighted (penalized) support volume the highest, maximal cross-sectional area the second highest, and part height the 3rd highest. The respondents put essentially no weight on the part-to-support surface area and parallelism with the recoater blade. Comparing weights (9) to Figure 8, you may think there is an error because the respondents did not keep the part height to be small.

Comparing weights (9) to Figure 8, you may think there is an error because the respondents did not keep the part height to be small. The team is still investigating this further, but this result should be correct, and may have to do with a correlation between V and A, a random nature of P, and an influence, but lack of correlation with other metrics, for H.Heuristics based optimization: ˆw can now be used to optimally orient any new part. First, a part is imported. Second, minimal and maximal values are found for metrics V, A, X, P, andH. Third, the objective function (8) is evaluated, yielding angles,   for printing. The,algorithm has been computation optimized to be very fast. The bar computes in 3 min and the handle computes in 3 hours, both on a 10 core cluster. Part orientations and support structures are provided in Figure 12.Figure 12. Optimal orientations and support structures using the survey-based heuristic optimization.3.2.

Part orientations and support structures are provided in Figure 12.Figure 12. Optimal orientations and support structures using the survey-based heuristic optimization.3.2. Novice Operators (Baseline Condition)Table 2 - Table 5 represent each novice operator’s attempt at the different parts in the baseline tests, with the highest attempt number being the successful print.BarSupport ModificationsPrint TimeSetupResult1540Attempt 1Attempt 2N/A23 hours 17 minutesStep in part at the location of different colorsN/A22 hours 48 minutesSuccessful printTable 2. Operator 1 rectangular bar result per attempt1541HandleAttempt 1Attempt 2BarAttempt 1Attempt 2Support ModificationsPrint TimeSetupResultN/A15 hours 49 minute sEdge delaminationHatching = 0.4 mm on far left side. Hatching = 0.8 mm on rest of part16 hours 1 minuteSuccessful PrintTable 3.

Operator 2 rectangular bar result per attemptSupport ModificationsPrint TimeSetup1543edge delaminationEdge delaminationEdge delaminationSuccessful PrintResultAttempt 1Border Teeth top length = 0.35 mm Hatching teeth top length = 0.7 mm Hatching 1.0 mm support angle = 45°=15 hours 40 minutesPhoto availableTable 5. Operator 2 contoured handle result per attemptAs shown in Table 2 - Table 5, each operator used support structure modifications to get a successful print. Once an orientation was chosen, it was not changed for subsequent attempts. Of the two test parts, the bar was the only one to have a different orientation between the operators. Operator one decided on a 45° print angle while operator two decided on a horizontal print orientation. Operator one believed that the flat part geometry would cause the bar to bow from distortion. The support structure in Attempt 1 was placed at both ends, thinking it would anchor the bar in place.

Operator one believed that the flat part geometry would cause the bar to bow from distortion. The support structure in Attempt 1 was placed at both ends, thinking it would anchor the bar in place. Upon seeing the result and the step in the part corresponding to the z-height of the support and part rejoining, operator one realized that the bowing would have already occurred before it reached this point and it wasn’t necessary. Operator one’s second attempt was to remove the tall support and leave just the base supported, resulting in a successful print.Operator two, in an attempt to reduce build time, orientated the part horizontally and increased the support structure hatch distance. This part failed to print as the support structure was not strong enough to withstand the deformation of the bar and jammed the machine. In the subsequent attempts, operator two utilized increasingly stronger support structure by reducing hatch distance and increasing top teeth length.

In the subsequent attempts, operator two utilized increasingly stronger support structure by reducing hatch distance and increasing top teeth length. This resulted in more connection points between the part and support, and ultimately, a successful print.Both operators chose the horizontal orientation for the handle and were able to apply the knowledge learned from the bar. Operator one’s first attempt was using the default support structure and resulted in part delamination at the edge. The logic used on operator one’s first bar attempt to prevent bowing and distortion was used on the second attempt of the handle and decreased the support hatch distance in the failed region. Again, this was an attempt to increase the connection points between the part and support, resulting in a successful attempt. Operator two utilized the same logic and increased the number of connection points between the part and support through support modification on their first attempt.1544not3.3.

Operator two utilized the same logic and increased the number of connection points between the part and support through support modification on their first attempt.1544not3.3. Expert survey based optimizationTable 6 illustrates the orientation recommendation from the expert survey based optimization.In the following sections of this report, these model recommended outputs were compared to the results achieved by the novice operators.Expert Survey OptimizationPrintSetupResultsTimeDidn’tCompleteHandleBarDidn’tCompleteTable 6. TCN-QTM Optimization model and Survey Optimization model recommended orientation and supports location for contoured handle and rectangular bar.The Expert Survey Optimization attempts on the handle and bar were both unsuccessful. In the case of the handle, the edge began to curl up causing a recoater arm jam. For the bar, the base support was not strong enough and failed during printing.

In the case of the handle, the edge began to curl up causing a recoater arm jam. For the bar, the base support was not strong enough and failed during printing. Upon recoating, the part was pushed further down into the powder, with successive layers printing on a bed of powder instead of the bar. This led to the stair case appearance on the part and poor sintering, as seen in the photo of Table 6, before ultimately causing a recoater jam. The failure of both parts of the Expert Survey Optimization could also be attributed to the low print angles (<35°).4. Discussion and ConclusionsExisting, commercial-grade support design software packages such as Magics™ have orientation design modules that use a weighting of different metrics (support area, height, etc.) to find an optimal orientation with respect to these weighted metrics.

However, it is up to the user1545to determine the weighting on each metric and the weighting values are not intuitive, nor have proper weightings been reported on. To remedy this lack of knowledge, the team set out to survey experts in the field of DMLS to better understand how they orient parts for DMLS. The survey was very unique compared to a simple multiple choice survey; users interacted with a CAD-like environment to orient parts as they best saw fit for DMLS printing. As DMLS is a nascent field, there are a limited number of people from which meaningful input can be extracted. Despite this complication, the survey had eighteen participants with a total of 151 different part orientations completed and the respondents spanned industry, national labs and agencies, and academia.

The analysis of the survey responses demonstrated that the DMLS researchers surveyed find orientations that maintain a small support volume, small maximal cross-sectional area of any given slice, and a small height in the z direction (see the weightings in Eqn. (9)). All three of these design objectives are intuitive. A small support volume leads to a shorter build time, less materials waste, and less support removal time. A small maximal cross-sectional area of a slice leads to less residual stress in the part from a smaller shrinkage in the x-y plane. A smaller height leads to less machine time because there will be fewer recoater blade passes. Interestingly, the survey respondents did not attempt to minimize the support-to-part surface area or the parallelism of slice edges with the recoater blade. Some users did comment that they did not think about the parallelism of the part edges with the recoater blade until they had made multiple submissions.

Some users did comment that they did not think about the parallelism of the part edges with the recoater blade until they had made multiple submissions. Although the team now has more information on how experts orient a part, they still failed to yield a favorable printing result with the optimization using a survey data based heuristic (Table 6). The rectangular bar and handle artifact prints failed at about mid build because the orientation algorithm selected an angle of the major axis of the part with in reference to the substrate that was too small. The team sees a single major issue why this may have happened. The .stl processor automatically generates support structures using the simple heuristic .stl. facets that have an angle between the facet normal and the substrate that is less than 35 deg should be supported. This is a common heuristic and used in many commercial software applications. However, this heuristic fails to account for edges in a part.

This is a common heuristic and used in many commercial software applications. However, this heuristic fails to account for edges in a part. Consider the simple part in Figure 13. The faces of the part are oriented such that the angle with the substrate is greater than 35 deg and hence should be unsupported. However, the line that makes the edge is less than 35 deg and should be supported. The team is currently working to fix this oversight in the code.Figure 13. Example of fault in support design heuristic. The bottom edge of the part is less than 35 deg from parallel with the substrate and should be supported. However, the intersecting1546faces that form the edge are both greater than 35 deg from parallel with the substrate, so the current heuristic will not support the faces.5. Acknowledgements and DisclaimerThis material is based on research sponsored by Air Force Research Laboratory under agreement number FA8650-12-2-7230. The U.S.

Acknowledgements and DisclaimerThis material is based on research sponsored by Air Force Research Laboratory under agreement number FA8650-12-2-7230. The U.S. Government is authorized to reproduce and distribute reprints for Governmental purposes notwithstanding any copyright notation thereon. The views and conclusions contained herein are those of the authors and should not be interpreted as necessarily representing the official policies or endorsements, either expressed or implied, of Air Force Research Laboratory or the U.S. Government.6. References[1] W. Frazier, “Metal Additive Manufacturing: A Review,” J. of Materi Eng and Perform, vol.23, no. 6, pp. 1917–1928, Jun. 2014.[2] “3D Printing in Aerospace - Revolution or Evolution? Overcoming Challenges for WiderAdoption,” Frost & Sullivan, 9831–22, Apr. 2015.[3] F. Calignano, “Design optimization of supports for overhanging structures in aluminum and titanium alloys by selective laser melting,” Materials & Design, vol.

PLOITING THE DESIGN FREEDOM OF RM D. M. Watts and R. J. Hague Rapid Manufacturing Research Group, Wolfson School of Mechanical & Manufacturing Engineering, Loughborough University, UKAbstract This paper details how Rapid Manufacturing (RM) can overcome the restrictions imposed by the inherent process limitations of conventional manufacturing techniques and become the enabling technology in fabricating optimal products. A new design methodology capable of exploiting RM’s increased design freedom is therefore needed. Inspired by natural world structures of trees and bones, a multi-objective, genetic algorithm based topology optimisation approach is presented. This combines multiple unit cell structures and varying volume fractions to create a heterogeneous part structure which exhibits a uniform stress distribution.Introduction Current product structures are not optimal; they are generally designed to meet a set of specifications identified in the conceptual design stage.

If optimisation is considered to be a process of improvement, then provided a design is within these acceptable design limits, it need not be improved upon further. Most industrial optimisation work is usually applied via a manually iterative approach often termed ‘over engineering’ which, although may improve a design’s performance, it does not yield a truly optimal solution as parts are often unnecessarily overweight (Mattheck 1998). At present, even if a truly optimal design is conceived it will still become compromised during the manufacturing stage due to the inherent limitations of conventional manufacturing technologies. For instance, injection moulding requires: the introduction of draft angles to allow the part to be removed from the tool; introduction of split lines; near constant wall thicknesses and the absence of overhangs within the design (Degarmo 1997).

In addition, further design considerations are required for assemblies to allow multiple parts to be assembled correctly, increasingly compromising a potentially optimal design. These Design for Manufacture and Assembly (DFMA) considerations (Boothroyd et al. 1994) effectively stifle designer’s freedom when creating new products and prohibit the production of globally optimal design solutions.1.Rapid Manufacturing (RM), due to its additive, layer-by-layer build nature, offers a notable alternative. RM allows the creation of virtually any given geometry without any need for tooling, allowing the fabrication of complex end-use parts that would be unmanufacturable via other means (Hopkinson et al. 2005). In addition, RM offers the ability to reduce part counts in assemblies by consolidating component parts into fewer pieces and, in some cases, can reduce the need for post production assembly operations by building components pre-assembled (Wohlers 2005).

Although still the subject of extensive research and development, RM is now a valid manufacturing method (Hopkinson et al. 2005).RM offers increased design freedom by removing the restrictions previously imposed by the DFMA considerations of conventional processes (Hopkinson et al. 2005). As a result optimal designs will have fewer compromises with RM becoming the enabling technology that will break the DFMA shackles of design. As suggested by the previous works of Campbell et al. (2003), Hague et al. (2003) and Mansour & Hague (2003), the advent of RM as a valid fabrication alternative compared to traditional processes will require a new methodology of Design for RM656(DFRM) to exploit this increased design freedom. It is the aim of this paper to propose a new design methodology that is capable of exploiting the increased design freedom offered by RM.

It is the aim of this paper to propose a new design methodology that is capable of exploiting the increased design freedom offered by RM. This is achieved initially by reviewing the optimisation found in the natural world followed by the inspection of numerical methods capable of simulating the design rules that Nature has followed in creating her optimal structures. This background is followed by the investigation of an existing freeware software (DesignLab) where the main topics of interest identified in the literature review are already combined, namely; Genetic Algorithms (GAs) and topology optimisation. The preliminary investigation highlights the benefits of a GA based topology optimisation system, but also points out the shortcomings of the particular software investigated, namely the inability to consider variable densities of a single material.

Finally, the results of this investigation are used to shape the proposed new methodology of truly optimal DFRM which is then presented. This novel GA based topology optimisation approach considers a single material of 3D unit cells, varying in volume fraction and form, to create a heterogeneous part structure. This method allows increasing part functionality through intelligent design towards an optimum.Literature Review Natural World Optimisation The Natural World has long been a source of inspiration for the problems that man faces with the field of Biomimetics dedicated to this technology transfer. Historical examples include the principal of flight, the fabric fastener Velcro® and the structure of the Eiffel Tower. More recent examples include Gecko tape, (Geim et al, 2003) and swimwear based on sharkskin, (Speedo 2005).2. 2.1Billions of years of evolution have led to the natural designs we see in the World today.

2.1Billions of years of evolution have led to the natural designs we see in the World today. Each natural design in existence is optimal for it’s specific environment, for if it were not, it would not have survived the test of time (Darwin 1859). Bones and trees are classic examples of optimal structures seen in the Natural World. Bones are able to adapt their microstructure to a changing loading environment over time, effectively adding more material where it is required and removing material where it is no longer needed. This process is known as Wolff’s Law, (Wolff 1892). Trees are also able to adapt their structures to changing loading environments as shown in numerous works by Mattheck (1998). Mattheck has shown that trees use their cambium growth layer to perform adaptive growth, quickly laying down additional material where high notch stresses occur and enabling shape adaptation to minimise the effects of prevailing winds (Mattheck 1990).

Unlike bones, however, trees are not able to remove material where it is no longer needed. Mattheck suggests this is due to bones belonging to animals that are required to move whereas trees remain stationary. Whatever the reasons for this behaviour, it is clear that Nature follows different design rules from mankind. If the optimal structures found in the natural World are considered to be the blueprints for optimal design, then the fundamental design rules that Nature follows, namely the efficient use of material and maintaining a uniform stress distribution, should be followed in engineering design.2.2Genetic Algorithms Nature’s process of survival of the fittest is a brutally efficient method of ensuring that only optimum design solutions survive. As such, the evolution process itself can be considered as an extremely powerful optimisation process and can be simulated in an accelerated form by the use of GAs.

As such, the evolution process itself can be considered as an extremely powerful optimisation process and can be simulated in an accelerated form by the use of GAs. GAs are general search and optimisation routines based on the mechanics of natural657selection and the Darwinian theory of evolution. Initially conceived by Holland (1975) and subsequently developed by the seminal work of Goldberg (1989), GAs harness the transfer of good genetic material from parent solutions to offspring solutions through the genetic operators of reproduction, crossover and mutation.The application of GAs as a means of component optimisation is usually performed in a parametric manner, where the GA is used to find the optimum values for all of the predefined component parameters (Renner & Ekárt 2003).

However, this type of optimisation does not allow for topological changes in geometry, such as the introduction of additional holes to reduce weight, (new parameters cannot be created), and as such only modifications to existing designs can be found instead of radical new design solutions. It is for this reason that a topology optimisation approach is needed.Topology Optimisation Nature’s design rule of only depositing material in structures where it is required with no excess to maintain a uniform stress distribution is the fundamental principle of topology optimisation; that is the economical distribution of material within a predefined design space. Hence topology optimisation is often referred to as the material distribution method. Topology optimisation is a numerical simulation technique that generates the optimal topological shape of a structure for a given mechanical load.

Topology optimisation is a numerical simulation technique that generates the optimal topological shape of a structure for a given mechanical load. The basic method solves the common engineering problem of determining the optimum arrangement of material distribution, with a limited amount of structurally isotropic material, within a given design space called the design domain.2.3In order for the material distribution and therefore the mass and associated structural behaviour to be represented adequately, the design domain is divided into many small discrete elements or cells. Loading and boundary conditions are applied by the user to the necessary discrete elements, thus completing the model.

Loading and boundary conditions are applied by the user to the necessary discrete elements, thus completing the model. Without any further guidance or decision making required from the user, the optimisation process systematically and iteratively eliminates and redistributes material throughout the design domain until a suitable structure is obtained (with regards to the objective function, constraints and termination criteria), thus yielding the final solution. Typically, this iterative process is implemented by either the addition or removal of elements from a Finite Element (FE) model.Many topology optimisation techniques are based on the Homogenisation Method, by Bendsøe & Kikuchi (1988), based on the use of an artificial composite material with microscopic voids using a variable density approach.

Instead of elements representing purely material being present or void of material, elements are allowed to have various densities (ranging between solid and void) throughout the optimisation process, yielding a solution in the form of a perforated composite material with a distributed microstructure. Indeed, it is the presence of the distributed microstructure with varying densities used in the homogenisation method that enables the resultant geometrical solutions to exhibit a uniform, and therefore optimal, stress distribution throughout the entire structure. From a conventional manufacturing point of view, the perforated regions within the composite structure representing the regions of intermediate density have little practical value and consequently these densities are usually forced towards either a 1 (representing solid) or a 0 (representing void of material) to define the final topology.

However, if RM is considered to be the chosen fabrication process, the occurrence of cells of intermediary658density is less of a problem and is something that can be exploited to increase part functionality and create a more uniform stress distribution.GAs can be applied directly to topology optimisation problems as shown by Kane & Schoeunauer (1996). The binary encoded string representation of the GA individual can be simply mapped into the design domain as a two dimensional (2D) array, as the cantilever beam example in Figure 1 shows. Here the solid and void regions are represented by the 1’s and 0’s of the binary code respectively.Figure 1 - Encoded Binary String Mapped into a 2D Cellular Array of a Cantilever BeamThis application of GAs to topology optimisation can also be extended to 3D problems by mapping the binary string into a 3D array.

When topology optimisation problems are solved by implementing a GA, the entire process will benefit from the nondeterministic search pattern that a stochastic process exhibits, therefore minimising the likelihood of premature convergence on a local optimum. This is arguably the opposite behaviour of methods that are variations of the deterministic homogenisation approach, where behaviour between successive iterations can be readily predicted and as such these methods are often trapped by local optima unless the number of discretised elements is high for a relatively small design domain.

By using a GA based topology optimisation approach, multiple materials can also be considered by moving away from a binary representation and introducing additional digits, as is shown in Figure 2 where 0’s represent void areas, 1’s represent material ‘A’ and 2’s represent material ‘B’.Figure 2 – Encoded String for Multiple MaterialsPreliminary Investigation Introduction As seen previously it is possible to combine a GA within the topology optimisation process, however, there is a lack of commercially available software packages capable of performing this function. One example that exists is a freeware, beta version software currently called ‘DesignLab 2005’ by DevDept (2005). This software was briefly investigated by the authors using a simply loaded cantilever beam problem.

This software was briefly investigated by the authors using a simply loaded cantilever beam problem. The software combines a GA and FE analysis with topology optimisation to create 3D components from either a single material or multiple materials by adopting a 3D array consisting of several different digits to represent the various materials.3. 3.1659DesignLab uses a basic GA that equates individual solution fitnesses by using a weighted sum of several objectives – weight, deflection, Factor of Safety (FOS), and introduces a penalty factor that reduces the combined fitness depending on how much an individual solution violates an objective. The fitness and penalty values are combined for each individual solution to calculate a cost or objective function. It should be noted that the cost function and associated cost values assigned in the software are merely a quantifiable measure of feasibility and have nothing to do with ecomomic or financial costing whatsoever.

It is currently impossible for the user to inspect or alter the weighting of the weighted sum in the beta version software.DesignLab was investigated to explore it’s potential as an RM design tool. For this investigation, a simple cantilever beam problem was considered where the design domain was divided by 20 cells in X, 10 cells in Y and 1 cell in Z with each cell having a length of side of 10mm. The design domain for this study was 3D but can be considered as 2½D being a simple extrusion in the Z-axis.

The design domain was fully constrained down the left hand side and a downward point load applied to the top right hand corner as can be seen in Figure 3.Figure 3 – Cantilever Beam Problem Defined in DesignLabThe parameters used to control the behaviour of the GA throughout this investigationwere the default parameters, namely:initial population size of 120 individuals with a population of 60 individuals maintained thereafter initial mutation probability rate of 0.015 dropping by 0.0025 for every 10 generations where results do not improve desired FOS of 2.25 with a penalty of 25 applied to individuals that violate this a minimum hole size of 1 cell(cid:131)(cid:131)(cid:131) (cid:131) (cid:131) mating by single point uniform crossover (cid:131)parent pairs selected by a weighted ranking of fitness valuesThese parameters were entered into the initial problem definition but were not altered atany stage of the investigation to ensure consistency.Initially only a single material was selected (Steel), allowing the GA to distribute solid material cells and void cells throughout the design domain.

This first scenario was repeated five times (trials A to E). The five beam structures of the first scenario were analysed to see if any cells were solid or void in all of the five trials. This was performed by superimposing the five structures on top of one another and adding them together. Cells that remained 0 were void in all five trials, cells that totalled 5 were Steel in all five trials and cells that were either 1, 2, 3 or 4660were Steel in some trials and void in others. The cells that remained solid or void throughout trials A to E were used to further constrain the design domain, thus steering the optimisation problem towards more desirable solutions based on the previous experience of the first scenario. The additional constraints were added to the cantilever beam problem definition creating a second scenario, which was also run five times (trials F to J).

The additional constraints were added to the cantilever beam problem definition creating a second scenario, which was also run five times (trials F to J). The five trials of the second scenario (F to J) were analysed in the same manner as those from the first, by superimposing the five beam structures on top of each other and adding them together to form the combined structure. This combined structure could be used to further constrain the beam problem design domain in additional scenarios.The effects on the overall stress distribution of introducing another material into the problem were of particular interest so the original cantilever beam problem was repeated with two materials, Steel and Aluminium. This third scenario was also run five times (trials I to M).The five beam structures of the third scenario, where two materials were considered, were also analysed in the same manner as before.

This third scenario was also run five times (trials I to M).The five beam structures of the third scenario, where two materials were considered, were also analysed in the same manner as before. However, this time, it was to see if any cells remained void, Steel or Aluminium in all of the five trials. The analysis was performed by altering the values of the cells for the different materials by a factor of ten so that Steel cells were represented by a 10, Aluminium cells were represented by a 1 and void cells were represented by a 0 as before. When the five structures were added together, cells that totalled 50 were Steel in all five trials, cells that totalled 5 were Aluminium in all five trials and cells that remained 0 were void in all five trials as before.

Cells that were anything other than 50, 5 or 0 in the combined structure fluctuated between Aluminium, Steel and void in the five trials.3.2Preliminary ResultsThe individual results of a single trial from the first scenario are shown in Figure 4, with both solid steel cells and void areas present. The stress distribution plot of Figure 4 shows that there are several areas of localised high and low stress; these have been circled in white. This solution can therefore be improved upon further to remove these localised stress concentrations and create a more uniform stress distribution.Figure 4 – Cantilever Beam Result from DesignLab for a Single Material, Showing Material Distribution Plot (a) and Stress Distribution Plot (b) with Areas of Local High and Low Stress Circled661Similarly, the individual results of a single trial from the third scenario can be seen in Figure 5, with solid Steel cells (dark grey), solid Aluminium cells (light grey) and void areas present.

The stress distribution plot of Figure 5 shows that there are fewer areas of localised high and low stresses when compared to the stress plot of the single material problem of Figure 4. This suggests that the stress distribution is more uniform and therefore the structure is more optimal than the single material structure of Figure 4.Figure 5 – Cantilever Beam Results From DesignLab for Two Materials, Showing Material Distribution Plot (a) and Stress Distribution Plot (b) with Areas of Localised Stress CircledDiscussion It was noticed that the average number of generations reduced from the first scenario to the second scenario. This can be explained by the presence of additional constraints imposed on the second scenario based on the analysed results of the first scenario.

This can be explained by the presence of additional constraints imposed on the second scenario based on the analysed results of the first scenario. These extra constraints effectively remove many structural permutations from the vast possibilities of the original problem, increasing the efficiency of the GA by preventing it from wasting time calculating already known poor solutions. Likewise, the average number of generations for the third scenario (two materials) is much greater than either of the single material scenarios due to the increased number of permutations possible for the structure. The GA now has to consider three options for each cell, namely void, Steel or Aluminium, as opposed to just the two of solid or void. By imposing additional constraints gathered from the analysis of the third scenario combined structure, it should be possible to reduce this average number of generations in subsequent scenarios.

By imposing additional constraints gathered from the analysis of the third scenario combined structure, it should be possible to reduce this average number of generations in subsequent scenarios. The introduction of additional constraints would once again reduce the number of permutations the GA may consider, thus increasing the efficiency of the GA in reaching a solution in a shorter time.4.The overall objective for these DesignLab optimisation problems is to minimise the cost function. The cost value is calculated by a weighted sum (as described earlier in Section 3.1) which is likely to include the weight of the solution structure, the maximum displacement (directly related to the stiffness or compliance of the structure) and the range of the FOS across the structure which bears a direct relationship to the stress distribution.

The average cost value from the first scenario to the second scenario was seen to improve and there was further improvement between the second and third scenarios. In each case the average FOS range has reduced resulting in a more uniform stress distribution across the structure (as was witnessed from Figure 4 to Figure 5). The average weight also reduced through all three scenarios, however, the maximum displacement reduced in the single material scenarios then increased in662the two material scenario. This is due to the inclusion of a less stiff material within the structure (Aluminium) which will deflect more readily than the stiffer Steel structures of the previous two scenarios.

This is due to the inclusion of a less stiff material within the structure (Aluminium) which will deflect more readily than the stiffer Steel structures of the previous two scenarios. The maximum displacement values may not be included in the weighted sum of objectives that form the cost function or, if they are included, do not have a large weighting as this increase does not seem to have a noticeable effect on the cost values.The most important observation in the results that should be noted is that the third scenario (two materials) had an improved average cost value and reduced FOS range when compared to either of the single material scenarios.

The stress plots of the structures in Figure 4 and Figure 5 show that the stress distribution is more uniform where multiple materials are considered compared to a single material with a homogenous structure, suggesting that future GA based topology optimisation systems should consider multiple materials in order to achieve a uniform stress distribution. The use of multiple materials within RM to increase part functionality has been considered by many researchers in the form of Functionally Graded Materials (FGMs), (Agarwala et al. 1992, Siu 2002, Tolochko 2003). However, there are many fabrication issues to be addressed in such cases in addition to the dilemma of recycling components fabricated of multiple materials.It is for these reasons that a single material only will be considered in the proposed methodology of optimal DFRM.

It should be noted that at this stage that although a single material is under consideration, the results from the first and second scenarios of the preliminary investigation (both single material only) suggest that a uniform stress distribution will not be achieved if cells are only allowed to be either 100% solid or 100% void. However, this phenomena is purely dependent on the size of the unit cells used compared to the size of the design domain. In order for any system to reach a truly globally optimal design solution, the geometry and topology should be completely free to ‘evolve’ into an ideal solution. However, this could only be achieved if the design domain is discretised into many small solid cells (voxels). This would have a considerable negative effect on required computation power, as a direct result of increasing the length of the GA bit string.

This would have a considerable negative effect on required computation power, as a direct result of increasing the length of the GA bit string. Despite the advances over recent decades in computing power, coupled with the decreasing cost of such equipment, the requirements of the idealised solution are still far beyond the resources of an average design engineer.The solution to the same cantilever beam problem obtained via the homogenisation method would yield a structure with a uniform stress distribution by permitting the density of the material to alter across the structure of the beam by varying the microstructure. The two material DesignLab solution of the same problem (shown previously in Figure 5) includes materials of considerably different densities, namely Steel and Aluminium, yielding a solution with a significantly more uniform stress distribution than any of the single material solutions.

In effect the Aluminium cells could be considered to be Steel cells of intermediary density. This phenomenon suggests that, in the absence of significant computing power to generate the voxel- sized elements (as outlined above), in order to achieve a uniform stress distribution across a solution structure consisting of a single material, variable densities of that material are a potential solution. This is only achievable when using an additive RM approach and must therefore be implemented into the new DFRM methodology.

This is only achievable when using an additive RM approach and must therefore be implemented into the new DFRM methodology. The DesignLab software has proved itself useful in these initial investigations, however, the ability to consider variable densities of a single663material is beyond it’s capabilities and as such the author is developing an alternative GA based topology optimisation software that is capable of this requirement.4.1Proposed Methodology of DFRMThe preliminary investigation of the DesignLab software in addition to the fabrication and recycling issues associated with Functionally Graded Materials has highlighted that a single material GA based topology optimisation system should be considered, where the uniform stress distribution is achieved by permitting variable densities of a single material.

If the density of a single elemental cell is considered to be the volume fraction of solid material within the whole unit cell, then this cell density or volume fraction can be varied by controlling certain feature parameters that describe a unit cell structure. Figure 6 shows a Computer Aided Design (CAD) model of a simple cell structure with various volume fractions achieved by altering the parameters of the hole features within the unit cell.Figure 6 – Varying Cell Volume Fraction by Alteration of CAD Parameter, in this Case the Hole DiameterIn addition to altering the volume fraction of a cell structure by controlling defining feature parameters within the 3D unit cell CAD model, it is also possible to create structures that have the same amount of material usage (identical cell volume fractions), that are capable of exhibiting different mechanical behaviour to differing loading conditions.

This suggests that cellular structures could be created that are suited to specific loading conditions, i.e. tailored specifically to the design needs. Some work in this field (mainly 2D) has been performed by Sigmund (1994) and Sigmund (1995). However, neither works considered mixing different unit cells within the same design domain, they merely solved the material distribution problem for an individual unit cell then arrayed that structure into a larger homogeneous assembly. In order for one structure’s mechanical behavioural response to be directly compared to another, they must have equal volume fractions. This will show which structures have the best distribution of material within each 3D cell.

This will show which structures have the best distribution of material within each 3D cell. Examples of various cell structures with a 50% volume fraction can be seen below in Figure 7.Figure 7 – Possible Cell Structures at 50% Volume Fraction664A topology optimisation process that uses a GA to create components consisting of a single cellular structure in varying densities as proposed should be able to create a cantilever beam similar to that shown in Figure 8a. The same approach, if considering multiple cellular structures, may create a cantilever beam component similar to that shown in Figure 8b. It must be noted that the proposed methodology of DFRM put forward is not a completely ideal solution to finding the globally optimal solution of a topology optimisation problem, as the optimisation will be limited to the various cell types that are defined in the initial problem.

As mentioned earlier, the ideal solution would be to use many tiny elemental cells but this is currently prohibited by the high computational power required by such as a system compared to typical computational resources available today. It is for this reason that there is a need for the proposed methodology of DFRM, as a current solution to creating optimised structures using readily available everyday computing resources to create the geometries and the evolving RM technologies as a means of fabricating these otherwise unmanufacturable designs.The proposed system will have three objectives. These are; firstly to minimise the mass of the geometry; secondly to minimise the maximum displacement (thereby increasing the stiffness); and thirdly to minimise the difference between the maximum and minimum Von Mises stress values. The third objective will be a measure of how uniform the stress distribution is throughout the structure.

The third objective will be a measure of how uniform the stress distribution is throughout the structure. In order to fully explore the trade offs and relationships that these objectives have on one another athey will not be combined into a weighted sum but will be all be measured equally using a Multi-Objective Genetic Algorithm (MOGA). The use of a MOGA will enable a Pareto set of equally optimal solutions to be generated and explored by the user.Figure 8 – Possible Cantilever Beam Geometry Created From (a) a Single Cellular Structure with Varying Densities and (b) Three Cellular Structures each with 50% Volume FractionsThe author is currently considering a simplified structural analysis where the test problems are simple load case scenarios with proven known results in order to validate the proposed methodology and refine the system parameters.

The optimal structures devised by Australian mathematician Michell (1904) will be used for this benchmark purpose in addition to simply loaded cantilever beam problems that should result in the solution structures depicted schematically in Figure 8.Conclusions The previous sections have highlighted how current product designs are not optimal due to the compromises that have to be made for the DFMA considerations of conventional techniques. Literature has shown that Nature follows a different set of design rules to man, namely the efficient use of materials and maintaining a uniform stress distribution. Nature’s design rules can be mimicked by implementing the topology optimisation method and the6.665evolution process itself can be simulated by a GA, both of which have been combined in the DesignLab software.

In the DesignLab investigation the stress distribution of the cantilever beams created using two materials was more uniform, and therefore more optimal, than either of the single material scenarios due to the presence of unit cells with differing mechanical properties. The homogenisation method relies on the creation of cells with varying densities in order to achieve a uniform stress distribution but this can cause manufacturing difficulties if fabrication is not via RM. The results of the preliminary investigation have showed that the DesignLab software is not capable of creating beam geometries with a uniform stress distribution unless multiple materials are considered. The ideal solution to a single material system would be to have many small solid elemental cells making up the design domain. This ideal system would be too computationally expensive as the GA bit strings would be extremely long.

This ideal system would be too computationally expensive as the GA bit strings would be extremely long. A new methodology and software system is therefore needed to allow these difficulties to be overcome. This has been presented as a current solution, and therefore a compromise from the ideal system mentioned above.The new design methodology for the creation of optimal structures that has been proposed uses a GA based topology optimisation approach based on the homogenisation method’s variable density approach. Part functionality is increased by simultaneously considering various cellular structures and densities to create a heterogeneous structure made from a single material.

Part functionality is increased by simultaneously considering various cellular structures and densities to create a heterogeneous structure made from a single material. This proposed methodology will be able to exploit the increased design freedom that RM technologies offer, enabling optimal design solutions with uniform stress distributions as seen in the natural World to be created that could not be manufactured via conventional techniques.Acknowledgements The authors gratefully acknowledge the financial support of the UK Engineering and Physical Sciences Research Council (EPSRC Grants IMCRC75 and IMCRC162).References Agarwala, M., Bourell, D., Beaman, J., Marcus, H., & Barlow, J., 1992, Direct Selective LaserSintering of Metals, Rapid Prototyping Journal, 1, pp. 26-36.Bendsøe, M.P., & Kikuchi, N., 1988, Generating Optimal Topologies in Structural Design Using a Homogenisation Method, Computer Methods in Applied Mechanics & Engineering, 71, pp.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEXPLORATION OF A CABLE-DRIVEN 3D PRINTER FOR CONCRETE TOWER STRUCTURESJesse J. Heineman*, Randall F. Lind*, Phillip C. Chesser*, Brian K. Post*, Alex M. Boulger*, Alex Roschli*, Lonnie J. Love*, Katherine T. Gaul*Oak Ridge National Laboratory, Manufacturing Demonstration Facility, 2350 Cherahala Blvd, Knoxville, TN 37932AbstractResearchers at Oak Ridge National Laboratory’s Manufacturing Research DemonstrationFacility (MDF) are currently developing a cable-driven concrete additive manufacturing (AM) system called SKYBAAM. This system is a novel solution for 3D printing large structures using concrete. The current research focuses primarily on proof of concepts for the cable driven system, material selection, material pumping solutions, and the concrete extruder design.

The current research focuses primarily on proof of concepts for the cable driven system, material selection, material pumping solutions, and the concrete extruder design. Looking forward from the success of the current research, this paper investigates the feasibility of using the SKYBAAM on a larger scale, specifically for extremely tall tower structures. The current system design presents challenges at a larger scale, and so the primary focus of this paper is to investigate new designs of a platform that would support large-scale SKYBAAM operations. Additionally, this paper will discuss the resulting deflections that can be expected due to machine operation and wind-loading.

Additionally, this paper will discuss the resulting deflections that can be expected due to machine operation and wind-loading. Excessive structural deflections could lead to loss of printing accuracy, or even a complete failure of the print, so it is important to establish that acceptable deflections can be reasonably achieved on these large-scale tower structures.IntroductionA novel solution for 3D printing large structures out of concrete is currently beingdeveloped at the MDF [1]. This cable-driven platform, called SKYBAAM, utilizes a mobile crane to support the extruder above the workspace, and four cable tension stations that are located on the ground to control the movements of the extruder in the X, Y, and Z directions.

The original vision for SKYBAAM is presented here (Figure 1), as well as the layout of a mid- scale prototype system that is currently being used at the MDF (Figure 2).1853-------------300ftlll- ----r-----------------►Figure 1: Potential Application for SKYBAAMApex supported by gantry craneFigure 2: Mid-Scale Prototype Platform at the MDFIn order to maintain stiffness in the X and Y directions and to keep the requisite crane loads to a reasonable level, the SKYBAAM architecture requires minimizing the cable angles from the cable tension stations to the extruder. However, minimizing this angle causes the overall footprint of the workspace to grow in the XY plane, and results in workspaces that are several times larger than the height of the structure being printed.

Scaling the current SKYBAAM platform up to a scale large enough to fabricate a tower structures in excess of 200 feet would require a very wide footprint, which would likely not be practical for most jobsites. In addition, a very large workspace footprint would necessarily require excessively long cables to connect the tension stations to the extruder. These longer cables would also need to be much thicker and maintain much higher tensions to keep the system stiffness to an acceptable level.1854The higher cable tensions would lead to higher loading requirements for the mobile crane supporting the extruder, as well as higher power requirements from the cable tension stations.

This paper will explore a new application for the SKYBAAM system that could be used to manufacture tower structures while reducing the workspace footprint, minimizing cable deflections, and reducing the cable station loading requirements compared to the traditional SKYBAAM arrangement.SKYBAAM for Tower StructuresThe basic design concept of SKYBAAM system has been preserved, with the primarydifference being the implementation of a superstructure that supports the weight of the extruder instead of a mobile crane. In addition, the cable tension stations would be mounted to the legs of the superstructure and would travel upward during a build instead of remaining stationary on the ground (Figure 3). This concept allows for a much smaller workspace footprint, which means that the cable lengths will be much shorter, which in turn will help minimize the cable deflections and tension station power requirements.

However, with this design, there will also be deflections of the tripod structure due to print head movements and wind loading. The primary focus of this paper is to ensure that we can reasonably achieve small enough structural deflections to maintain high print accuracy with this type of superstructure.The primary superstructure design that was investigated was a tripod-like structure, with each of the legs made from a square-profile truss beam. Two different types of tripod structures were analyzed. In each case, the dimensions of the leg truss structure and the sizes of the main load-carrying elements were optimized to achieve roughly 1/8 in. deflection due to the variable print head loads and roughly ½ in. deflection due to an evenly distributed 30mph wind load.

deflection due to the variable print head loads and roughly ½ in. deflection due to an evenly distributed 30mph wind load. All the following structural analysis was performed in SOLIDWORKS simulation.Figure 3: Tension Station ArrangementAngled tripod legsThe first type of tripod structure that was analyzed was made up of angled tripod legs thatare spread out at the base and meet at a single point at the apex (Figure 4). The resulting optimized truss size for this design was a 72” square profile made up of 4” x 3.5” round tubes as1855the main structural elements. The overall weight for this type of structure is approximately 100,000lb.Figure 4: Angled Leg Tripod StructureDeflection due to gravityThis type of design requires long truss legs to achieve the required clearance above theprinted part due to the angle of the truss legs.

In addition, due to the angle of the legs, there is a component of gravity acting perpendicular to the truss leg, which causes a static deflection as shown in Figure 5.1856URES(in)6.148 e-01S.636e-01S.124e-01_ 4.611e-014.099t-013.074e-012.049e-011.02Se-01S.124e-023.937e-32Figure 5: Deflection of Angled Leg Truss Due to GravityDeflection due to combined loadsThe maximum load acting on a truss leg is composed of gravity, the acceleration of the print head, and tension from the opposing tension stations. The geometry of each of the loads was analyzed to determine the components that act perpendicular to the truss leg, which were used for the simulation. The print head loads were applied mid-way between supports to simulate the maximum deflection situation.

The print head loads were applied mid-way between supports to simulate the maximum deflection situation. The deflection is shown in Figure 5.URES(in)6.689'!,-016.00le-01_ S.473e-014.86Se-012.432e-011.824e-011.216e-01Figure 6: Deflection of Angled Truss Leg Due to Combined LoadsDeflection due to print head loadsThe static components of the deflection can be easily predicted and managed, so it is thevarying loads due to the print head movements that we are primarily interested in. These movements can be variable during a build cycle and are not as easy to manage. The variable deflections can be calculated by subtracting the static gravitational forces from the total deflection due to the combined loading.

The variable deflections can be calculated by subtracting the static gravitational forces from the total deflection due to the combined loading. The maximum deflection for this structure is shown to be within desired range of approximately 1/8” (0.73”[Figure 5] – 0.61”[Figure 6] = 0.12”).Deflection of angled leg tripod structure due to wind loadsDeflections from the wind loading simulation are shown in Figure 7, with the maximumdeflection of 0.36” occurring at the top of the tower. A more meaningful deflection location1857would be at approximately 3/4 height, where the maximum build height would occur.

A more meaningful deflection location1857would be at approximately 3/4 height, where the maximum build height would occur. This location has a deflection closer to 1/4", which is within the desired range of 1/2”.URES(in)3,622 e-013.320e-01r 3,01~-01_ 2.717e-01_ 2.41Se-012.113 e-011.811 e-011.50~-011.207 e-019.056e-026.037 e-023.01~-023,937 e-32Figure 7: Deflection of Angled Leg Tripod Structure Due to Wind LoadVertical tripod legsThe second type of tripod structure that was analyzed included a large roof structure at the top, which allows the truss legs to be shorter and completely vertical as shown in Figure 8. The optimized truss size for this design was a 36” square profile made up of 3” x 2.5” round tubes as the main structural elements.

The optimized truss size for this design was a 36” square profile made up of 3” x 2.5” round tubes as the main structural elements. The overall weight for this type of structure is approximately 60,000lb.1858Figure 8: Vertical Leg Tripod StructureDeflection due to print head loadsThe print head loads were applied mid-way between supports to simulate the maximumdeflection situation (Figure 9). The maximum deflection is shown to be 0.13”, which is very close to the target value of 1/8”.URES (in)1.302e-011.193e-011.C65e-01_ 9.761e-02_ 8.677e-027.592e-026.5C6e-025.423e-024.338e-023.254e-022,169e-021.C65e-023.937e-32Figure 9: Deflection of Vertical Leg Truss Structure Due to Print Head LoadsDeflection due to wind loadDeflections from the wind loading simulation are shown in Figure 10, with the maximumdeflection of 0.54” occurring at the top of the tower.

A more meaningful deflection location would be slightly below the top of the tower, where the maximum build height would occur. This location has a deflection closer to 0.45”, which is within the desired range of 1/2”.1859URES (in)S.438e-014.9:JSe-Ol_ 4.S32e-01. 4.079e-Ol3.625e-Ol3.172e-012.7l9e-012.266e-01l.8l3e-Ol_ l.31:0e-019.063e-024.S32e-023.937e-32Figure 10: Deflection of Vertical Leg Tripod Structure Due to Wind LoadConclusionBoth types of structures have achieved our target deflection values for both print headand wind loads. The angled tripod structure has fewer components and a simpler assembly which could have a relatively lower installation time and cost. However, it would require a taller and higher capacity crane. On the other hand, the vertical tripod structure has a significantly lower weight but requires more tensioning elements between the truss legs, which could increase the installation time and cost.

On the other hand, the vertical tripod structure has a significantly lower weight but requires more tensioning elements between the truss legs, which could increase the installation time and cost. Further analysis should be performed to compare the installation costs of each type of structure, as well as any other jobsite logistics to help guide the future design of these structures.References[1] Chesser, P., Post, B., Roschli, A., Lind, R., Boulger, A., Love, L., Gaul, K., FieldablePlatform for Large-Scale Deposition of Concrete Structures. Solid Freeform Fabrication Symposium. Austin, TX, 2018.1860.

 EXPLORATION OF BINDER JETTING OF COPPERYun Bai and Christopher B. WilliamsDesign, Research, and Education for Additive Manufacturing Systems LaboratoryDepartment of Mechanical EngineeringVirginia Polytechnic Institute and State UniversityAbstractThe ability to fabricate geometrically complex copper shapes via Additive Manufacturing (AM) could have a significant impact on the design and performance of thermal management systems and structural electronics. In this research a Binder Jetting AM process (ExOne R2) was used to fabricate green parts made of high purity copper powder. Once printed, the green part was sintered under a reducing atmosphere to create copper parts in pure metal form. The authors varied (i) powder size, (ii) sintering profiles, and (iii) atmospheric control to explore their effects on final part density and shrinkage. The sintered part density was 85% of the theoretical value due to the relatively coarse powder and loose packing of the powder bed.

The sintered part density was 85% of the theoretical value due to the relatively coarse powder and loose packing of the powder bed. The result demonstrates the feasibility of using Binder Jetting to create copper parts with complex geometries.1. Introduction1.1 Motivation for the Additive Manufacturing of copperthe 2009 Roadmap for Additive Manufacturing was to use Additive Manufacturing (AM) to create complex heat exchangers to enable a new generation of power generators for portable electronic devices that use hydrocarbon fuel cells [1]. Hydrocarbon fuels have higher energy density than batteries, but one challenge facing the use of hydrocarbon fuel in electronic devices lays in the thermal management where heat loss needs to be minimized.

Copper, a highly conductive material, could be used to fabricate the reactor, which would feature complex intake and exhaust passages that can minimize heat losses via heat recirculation.One of the research opportunities identified inHowever, advances in the design of highly efficient thermal management systems are somewhat stymied by an inability to Additively Manufacture complex structures with copper material. To release this design constraint, the authors are investigating the use of Binder Jetting to process copper. This layer by layer fabrication process offers the utmost design freedom in the realization of complex geometries.1.2 Current copper Additive Manufacturing technologiesPure copper material has been fabricated with Additive Manufacturing processes including Laser Sintering (LS), Electron Beam Melting (EBM) and Ultrasonic Object Consolidation (UOC).Laser Sintering (LS) has been used to indirectly fabricate metal parts from metal-polymer powders [2,3].

In their work, Badrinarayan and Barlow used LS to fabricate copper parts by selectively laser-scanning a bed of a copper-polymer (PMMA) powder mixture to melt the polymer layer-by-layer and create copper green parts. The copper green parts are then fired at elevated temperature to burn off the polymer, which acted as an intermediate binder, and sintered in a reducing atmosphere. This process was shown to create copper parts with density of 48% of the theoretical value, and up to 60% if using bimodal powders.Copper components and open-cellular copper structures have been successfully created via Electron Beam Melting (EBM) from atomized copper precursor powders [4–6]. Copper oxide precipitates, and a novel columnar architecture by the reorganization of these precipitates, was observed due to the affinity of oxygen of finely atomized copper.

Copper oxide precipitates, and a novel columnar architecture by the reorganization of these precipitates, was observed due to the affinity of oxygen of finely atomized copper. In another study of processing pure copper parts via EBM, copper lattice structures and a near fully dense copper parts (99.8% of the theoretical density) were achieved via EBM [7,8]. The authors noted that the resultant parts suffered from dimensional inaccuracy (more than 13%) and contained pores and bubbles. These defects are believed to be caused by the severe dissipation of thermal energy during the melting process due to the high thermal conductivity of copper.Ultrasonic Object Consolidation (UOC), a laminate-based AM process that features stacking, joining, and machining 2D laminates, has been used to fabricate copper parts.

Tapes of copper were successively stacked and joined via ultrasonic welding, in which the recrystallized fine grains were converted from coarse-grained structures [9–11].1.3 Rationale of Binder Jetting of copperIn this paper the authors explore the use of Binder Jetting to create copper parts. Binder Jetting is an Additive Manufacturing process in which a liquid binding agent is selectively deposited to join powder materials (Figure 1) [12]. In the green part creation stage, an inkjet printhead selectively deposits binder droplets into a powder bed and the binder interacts with the powder particles to form primitives that stitch together to form a cross-sectional layer. Once a layer has been printed, the powder feed piston raises, the build piston lowers, and a counter- rotating roller spreads a new layer of powder on top of the previous layer. The subsequent layer is then printed and is stitched to the previous layer by the jetted binder.

The subsequent layer is then printed and is stitched to the previous layer by the jetted binder. The remaining loose powder in the bed supports overhanging structures and is removed with compressed air in post- processing. The green parts are then heated at an elevated temperature to cure the binder, which gives the parts sufficient strength for cleaning and handling. The printed parts require sintering to obtain the final density and strength. During sintering, the binder burns off and then particles sinter together through atomic diffusion.Figure 1. Binder jetting schematic [13]Binder Jetting has several characteristics that make it well-suited for fabricating complexcopper geometries:Its use of a powder bed eliminates the need for support structures typically required creating for overhanging features. While other direct-metal AM systems employ a powder bed, they require additional structures for part anchoring or heat dissipation purposes.

While other direct-metal AM systems employ a powder bed, they require additional structures for part anchoring or heat dissipation purposes. Since Binder Jetting does not require such structures, this effectively eliminates the effort of printing and removing support structures.Binder Jetting is readily adaptable to a wide-range of materials [14]. Material development for direct-metal AM systems is typically hindered by concerns over a material’s melt point, reactivity, optical reflectivity, and thermal conductivity. A variety of material systems, including polymer [15–19], ceramic materials [20], metals [21], and foundry sand [13,22] have been adapted to Binder Jetting.As it does not require an enclosed chamber, and does not use expensive energy sources, Binder Jetting is an inherently scalable technology.

Binder Jetting systems have large build volumes; for example, ExOne offers build volumes of up to 780 x 400 x 400 mm for metal powders [36]; VoxelJet offers a “continuous” sand printer that does not restrict the length of its prints (at a 850 x 500 mm print width and height [37]). These large part sizes are possible as Binder Jetting is free from the powder bed thermal management constraints typically found in direct-metal AM processes. Furthermore, Binder Jetting systems have a relatively high throughput: a 100 nozzle printhead can create parts at up to ~200 /min.As Binder Jetting of metal functionally separates part creation from powder sintering, one can leverage understanding gained from well-studied traditional powder metallurgy (P/M) processes.1.4 ContextIn this paper, the authors present the results of their preliminary exploration of processing copper via Binder Jetting.

Green parts were first fabricated by printing three different atomized copper powders (median size ranging from 15 to 75 µm) with an ExOne R2 3D printer. The green parts were then sintered in a tube furnace featuring a controlled atmosphere. The experimental procedure (Section 2) followed an established process for developing new materials for Binder Jetting, which included experiments in binder selection, powder formulation, powder-binder interaction testing and printing, and post-processing [14,23]. The resultant copper parts were characterized for sintered density, shrinkage, porosity, chemical composition, and ultimate tensile strength (Section 3). A summary of the work is presented inSection 4 along with a discussion on future work on how to overcome the challenges encountered in Binder Jetting copper.2.

A summary of the work is presented inSection 4 along with a discussion on future work on how to overcome the challenges encountered in Binder Jetting copper.2. Experimental procedureGiven the authors’ goal of processing copper via Binder Jetting, they followed Utela and coauthors’ [23] established methodology for developing new materials for Binder Jetting. The methodology includes the selection of binder and powder, the determination of printing parameters in order to fabricate green parts, and the determination of the appropriate post- process sintering cycle. In this section, the specific experimental procedure for each stage of the development process is presented.2.1 Binder selectionThe criteria for binder selection for Binder Jetting focused on (i) binder interaction with candidate powders (wettability and penetration) and (ii) binder residue in the debinding post- process.

The authors chose to work with ExOne’s standard binder (PM-B-SR-1-04), as it is widely used binding agent for many metal powders and has been shown to have minimal ash residue during debinding.The functional ingredient of the binder is a thermosetting polymer that can react to form polyester resins, which harden upon heating. The printed parts are cured at 190 °C for two hours to fully set the binder and provide satisfactory green part strength. The TGA analysis has demonstrated the binder can completely burn off at 450 °C.2.2 Powder selectionPowders are selected based on their particle size distribution, morphology and chemical composition. In P/M, fine powder is preferred in order to lower the required sintering temperature and to improve densification. However, in Binder Jetting, particles larger than 20 µm are typically preferred so that the powder can be successfully spread during the recoating step [24].

However, in Binder Jetting, particles larger than 20 µm are typically preferred so that the powder can be successfully spread during the recoating step [24]. Small particles can be used, however they need to be controlled in a small volume percentage and generally cannot be smaller than 1 µm [23,24,34]. A spherical particle shape is preferred over irregular shape because it tends to flow during recoating and it also is more easily wetted with binders.Given these considerations, the authors chose gas atomized copper powder that features spherical particle shapes for Binder Jetting of copper. Three different powders were explored in order to determine the effect of powder size distribution on part processing: AcuPowder 153A copper powder, AcuPowder 500A ultra-fine atomized copper powder, and Ozometal atomized copper powder.The particle size distribution of received powders was analyzed by laser scattering (Horiba LA-950 Laser Scattering Particle Size Distribution Analyzer).

A FEI Quanta 600 FEG Environmental Scanning Electron Microscope (ESEM) using a voltage of 20 kV revealed the particle shape and verified particle size information by taking microscopic pictures of discreteparticles. The chemical composition was determined by Energy-dispersive X-ray spectroscopy (Bruker EDX) using the same voltage in ESEM.The strength of the green part, and ultimately the quality of the final sintered part, is determined primarily by the green part density. This density is directly dependent on the density of the powder bed of the machine, which is affected both by powder characteristics and recoating parameters (e.g., the rotational and translational velocities of the counter-rotating spreading mechanism). To evaluate the powder bed packing density, the authors printed a density measuring cup into the powder bed, as shown in Figure 2. The mass of the packed powder that was in the cup was then measured.

To evaluate the powder bed packing density, the authors printed a density measuring cup into the powder bed, as shown in Figure 2. The mass of the packed powder that was in the cup was then measured. The packing density was then calculated as the ratio of the mass of the poured powder and the known volume of the density measuring cup. This method was also used to estimate the packing density for calculating the binder saturation (section 2.3).(a) (b) (c)Figure 2. Powder bed packing density measurement method: (a) CAD file of the density measuring cup with a known volume. (b) The printed density measuring cup with packed powder. (c) The packed powder was emptied from the cup to measure the packing density2.3 Green part creationThe compatibility of the received powder with the ExOne metal binder was tested by a bench top powder-liquid interaction test before the printing [23]. A pile of powder was strafed with a syringe of the binder liquid.

A pile of powder was strafed with a syringe of the binder liquid. The absorption rate of binder and the effectiveness of bounding particles were examined.The critical printing parameters in the green body creation include layer thickness and binder saturation ratio. For a successful recoating of powders in Binder Jetting, the layer thickness should be at least thicker than the largest particle and preferably three times of particle size [25]. Saturation ratio (S) describes the percentage of the air space between powder particles that is occupied by a binder volume ( ). The saturation ratio is determined from Equations (1) and (2), and is based on the packing density of the powder bed (PR) and the volume of solid particles in a defined envelope ( ).

The saturation ratio is determined from Equations (1) and (2), and is based on the packing density of the powder bed (PR) and the volume of solid particles in a defined envelope ( ). From these measurements, the ExOne R2 printing software calculates the required volume of binder to be printed once the binder saturation is input by the user.⁄(1) (2)Saturation needs to be carefully selected as it affects the quality of printed green parts as well as the final sintered density. Sufficient binder saturation is required to obtain green part strength.Conversely, oversaturation of the powder results in dimensional inaccuracy and low sintered density because of bonding unwanted loose powders and the resulting porosity following binder burn-out.

For spherical particles, the binder saturation is relatively low because it uses binders efficiently by having small necking area between particles [26].The layer thickness and the binder saturation of each copper powder were determined experimentally based on these criteria. The binder saturation was estimated first based on the observation from the powder-binder interaction test. The specimens with different combinations of layer thickness and binder saturation were printed iteratively until satisfactory green part quality and strength were observed.2.4 Post-processingTo determine the suitable sintering profile for each powder, and to investigate how sintered densities are affected by sintering temperature, time and atmosphere, rectangular sintering test coupons (24 × 8 × 4 mm) were printed from each powder.

The density and dimensions of each coupon were measured before and after the sintering to evaluate densification and shrinkage.The solid state sintering of copper was conducted in a MTI GSL-1600X tube furnace (101 mm tube O.D.) in either air with a vacuum or in a controlled hydrogen/argon atmosphere as shown in Table 1.Table 1. The atmospheres used for sintering copperAtmosphere Hydrogen/Argon AirFlow rate (ml/min) or vacuum (torr) 100/0, 80/20, 60/40, and 20/80 10-200 torrWhen post-processing the parts in an air/vacuum environment, the heating schedule consisted of two primary stages: debinding and sintering. A 30 minute hold occurred at 450 °C to ensure removal of the printed binder. The temperature and time of the sintering stage (Figure 3a) was varied in order to explore its effect on sintered part density. Heating ramps of 5 °C /min and a cooling ramp of 5 °C /min were used to transition between stages.

Heating ramps of 5 °C /min and a cooling ramp of 5 °C /min were used to transition between stages. When post-processing in a reducing environment (Section 3.3), the heating stage featured three steps (Figure 3b). The additional step was held at 700 °C to facilitate the reduction of copper oxides before the closure of pores.(a) (b)Figure 3. The heating schedule used for sintering copper (a) in air/vacuum and (b) in reducing atmosphere2.5 Part characterizationThe densification of copper was evaluated by comparing powder packing density, green density and final density. The green and sintered density was measured by a helium gas pycnometer (micromeritics AccuPyc II 1340) and a density measurement kit using Archimedes’ principal (Denver Instrument). The sample mass was measured by an analytical balance (ACCULAB AL-64). Helium gas pycnometry measures the total displaced skeletal volume of the sample [27].

The sample mass was measured by an analytical balance (ACCULAB AL-64). Helium gas pycnometry measures the total displaced skeletal volume of the sample [27]. Archimedes’ principle measures the volume of water displaced by an immersed object, of which the surface pores were sealed by oil impregnation before measurements [28]. The volume was also estimated by measuring linear dimensions of the rectangular bar for reference purpose. The measured density was converted to the relative density by comparing with the theoretical density of the pure copper (8.96 g· ) at room temperature.Shrinkage is produced by the diffusion mechanisms including viscous flow, volume diffusion and plastic flow [29]. The shrinkage of sintered copper samples was evaluated as linear shrinkage and volumetric shrinkage by measuring the dimensions of the sample before and after sintering.The microstructural features of the sintered samples were examined by microscopy.

Magnified images were generated at lower magnification by reflected light microscopes (Olympus Transmitted/reflected Light Microscopes). Higher magnification images were generated via electrons by ESEM (FEI Quanta 600 FEG Environmental Scanning Electron Microscope) using an accelerating voltage of 20 kV. Weight percentage of Cu and O in the green and sintered copper samples was determined by EDX (Bruker Energy-dispersive X-ray spectroscopy).To evaluate the strength of the final parts, tensile test specimens were printed with AcuPowder 500A copper powder and sintered in hydrogen/argon at various temperatures. The tensile test specimen was designed following “Standard Flat Unmachined Tension Test Specimens for Powder Metallurgy (P/M) Products” in ASTM E8 standard [30]. The printed dimension was scaled up by a scaling factor based on the average shrinkage detected in the similar sintering profile.

The printed dimension was scaled up by a scaling factor based on the average shrinkage detected in the similar sintering profile. The stress-displacement curve was obtained from an Instron (150 kN static rating) testing machine with a displacement rate of 50 mm/min.3 Results and Discussion3.1 Green part creationTable 2 summarizes the powder information used in creating copper parts via Binder Jetting.The information on particle size distribution, weight percentage of copper element and the packing density was determined by the characterization techniques presented in Section 2.2. The particle size distribution and the powder morphology SEM pictures are shown in Figure 4 andFigure 5. The Ozometal atomized copper powder has the highest packing density because of the relatively wide particle size distribution.Table 2.

Particle size distribution determined by laser scattering of (a) AcuPowder 153A copper powder (b) Ozometal atomized copper powder (c) AcuPowder 500A atomized copper powder(a)(b)(c)Figure 5. Powder SEM microscopy of (a) AcuPowder 153A copper powder (b) Ozometal atomized copper powder (c) AcuPowder 500A atomized copper powderIn a bench-top powder-binder interaction test where binder droplets were dispensed from a syringe onto a pile of powder, the binder was demonstrated to have a good wetting and penetrating characteristic with the atomized copper powders. Rectangular test specimens were successfully created by Binder Jetting on ExOne R2 3D printer with the settings listed in Table 3. The layer thickness and binder saturation were determined using the methods described in Section 2.3.

The layer thickness and binder saturation were determined using the methods described in Section 2.3. The layer thickness was adjusted based on the median/largest particle size, and a binder saturation of 70% was demonstrated to have a good shape creation as well green part strength for all three powders. Two binder saturation ratios, 60% and 80%, were used for creating AcuPowder 500A copper powder to study the effect of binder saturation on sintering characteristics (Figure 10).Table 3.

Two binder saturation ratios, 60% and 80%, were used for creating AcuPowder 500A copper powder to study the effect of binder saturation on sintering characteristics (Figure 10).Table 3. Printing parameters in the green part creationPowder nameLayer thickness (µm)Binder saturation ratio (%)AcuPowder 153A copper powder Ozometal atomized copper powder AcuPowder 500A ultra-fine atomized copper powder1008080707060 and 803.2 Preliminary study on the sintering of copper printed via Binder JettingTo identify the appropriate sintering profile for the printed copper green parts, the authors conducted a preliminary sintering study that leveraged post-processing procedures used in existing copper powder metallurgy research. In a sintering study of compacted copper powder, satisfactory densification was achieved in a temperature range from 750 °C to 1000 °C, and a maximum of 90% sintered density was observed when sintered at 1000 °C for 100 minutes [35].

In the sintering of copper powder in injection molding (MIM) fine oxide reduced copper with a median size of 10 µm was sintered below 1050 °C where a sintered density of 95% was achieved [32].Figure 6. Sintered copper part samples (AcuPowder 153A copper powder sintered in pure hydrogen at 1000 °C for 8 hours)Figure 7. SEM (left) and optical microscopic (right) pictures showing the necks between particles formed in the initial sintering stageWhen these same post-processing parameters were used on copper powder processed by Binder Jetting, the produced parts had significantly lower density. Figure 6 shows the samples made from AcuPowder 153A copper powder (74 median size) being fired at 1000 °C for 8 hours. The shrinkage after sintering was less than 5% and the sintered density was less than 60%. The SEM picture in Figure 7 shows the discrete particles and the sharply concave necks formed between particles.

The SEM picture in Figure 7 shows the discrete particles and the sharply concave necks formed between particles. This is the evidence of an initial stage sintering where no significant densification effect and the recrystallization of the structure and pores was observed.The difficulties in sintering copper prepared by Binder Jetting are not surprising given the large powder particle size and loose packing of the particles. These characteristics leave large spaces between particles which lowers the sinterability of solid state sintering. Similar results were found in prior work in using LS to process copper. In that work, the researchers looked to bimodal mixing of powders and polymer-coated powders to increase part density [2].

Similar results were found in prior work in using LS to process copper. In that work, the researchers looked to bimodal mixing of powders and polymer-coated powders to increase part density [2]. In this paper, the authors look to achieve higher part density by exploring different heating schedules and sintering atmospheres, and by using powders with various particle size distributions.3.2 Densification and shrinkageDue to the unsuccessful densification effort mentioned in the previous section, higher sintering temperatures were examined. Samples made from AcuPowder 153A copper powder and Ozometal atomized copper powder were sintered in air/vacuum air at temperatures ranging from 1060 °C to 1090 °C to explore its effect on volumetric shrinkage (Figure 8b) and part density (Figure 8a).

Ozometal copper powder can be sintered to a significantly higher density than AcuPowder 500A due to the smaller size and wider distribution.10060) l a c i t e r o e h t%(y t i s n e D d e r e t n i S9080706050)%( e g a k n i r h S c i r t e m u o Vl5040302010401050106010701080109011000 10501060107010801090Sintering Temperature (°C)Sintering Temperature (°C)AcuPowder 153AOzometal powderAcuPowder 153AOzometal powder(a)(b)Figure 8. The relationship between densification and sintering temperature of AcuPowder 153A and Ozometal powder: (a) sintered density (b) volumetric shrinkage, sintered in air/vacuum for 4 hrs.Finer copper powders, Ozometal copper powder and AcuPowder 500A, were sintered in hydrogen/argon atmosphere. The sintering temperature (Figure 9) and time (Figure 10) were varied to study their effects on the sintered density and shrinkage.

The sintering temperature (Figure 9) and time (Figure 10) were varied to study their effects on the sintered density and shrinkage. In Figure 10, two different binder saturation ratios were sintered using same heating schedules.Higher sintering temperature increased the final sintered density and shrinkage but was limited by the powder size and packing. Longer sintering time has the tendency to improve the densification. However in Figure 10 the best sintered density (85.5%) occurred at 4 hours instead of 6 hours. This observation needs further investigation by expanding to different powders and wider range of the sintering time. In Figure 10 the binder saturation did not have a major effect on densification; however more findings with respect to the binder saturation are discussed in Section 3.3.

In Figure 10 the binder saturation did not have a major effect on densification; however more findings with respect to the binder saturation are discussed in Section 3.3. While it can be concluded that small particle size improves sinterability and sintered density, it is noted that the particle size distribution contributes to sinterability, as the wide size distribution in Ozometal atomized copper powder sintered to similar density of the finer powder.11008540) l a c i t e r o e h tf o %(y t i s n e D d e r e t n i S807570656055)%( e g a k n i r h S c i r t e m u o Vl3530252015105501030 1040 1050 1060 1070 1080 10900 1030 1040 1050 1060 1070 1080 1090Temperature (°C)Temperature (°C)AcuPowder 500AOzometal PowderAcuPowder 500AOzometal Powder(a)(b)Figure 9.

It was noticed the copper samples being fired at high temperatures that are close to, or even higher than, the melting temperature of copper did not melt or lose the shape. This is caused by the powder impurity or copper oxidation layers, which can increase the powder melting temperature.Table 4. The highest achievable sintered density and volumetric shrinkagePowder nameSintered profileSintered density (% theoretical)Volumetric shrinkage (%)AcuPowder 153A copper powder1090 °C/240 min. Air/vacuum63.2%14.9%Ozometal atomized copper powder1080 °C/120 min. hydrogen/argon77.6%22.7%AcuPowder 500A ultra-fine atomized copper powder1080 °C/240 min. hydrogen/argon85.5%43.4%For AcuPowder 153A copper powder, which is the coarsest powder among the three receivedpowders (Table 1), the density after sintering increased from 55% to less than 65% and the volumetric shrinkage was less than 15%.

The sintered density of a finer powder (Ozometal atomized copper powder) improved by 10% compared to the coarse powder, which reached an average of 74% of the pore-free copper density. The maximum sintered density of AcuPowder 500A ultra-fine atomized copper powder (the finest among three) was 85.5%.3.3 Porosity typeThree different density measurement methods were used to determine sintered density. Thevolume calculated from the three dimensions of rectangular bars is a rough estimation for the bulk volume. The displacement volume measured by Archimedes’ principal and gas pycnometry represents the bulk density only if the surface porosity is sealed by oil impregnation. For non- impregnated specimens, the bulk volume does not include the surface connected pores, which are filled with water or helium gas.

For non- impregnated specimens, the bulk volume does not include the surface connected pores, which are filled with water or helium gas. In this paper the Archimedes’ principle density of oil impregnated parts is used to measure the bulk density.The authors noticed non-impregnated sintered parts have a nearly theoretical density of copper when measured by gas pycnometry. This result indicates that the majority of the pores in the sintered copper samples are surface-connected. This conclusion is also supported by examining the porosity of sintered copper under SEM (Figure 11).In Binder Jetting, one droplet of binder jets into the powder bed and binds multiple particles together forming a primitive. In Figure 11 (a) and (b), one can tell the surface of the sintered part is composed of vertical lines (Y-direction) formed by stitching primitives together.

In Figure 11 (a) and (b), one can tell the surface of the sintered part is composed of vertical lines (Y-direction) formed by stitching primitives together. In Figure 11 (c), the copper particles in the same primitive sintered to final stages where pores inside one primitive are very less. However, the diffusion and necking between adjacent primitives are very limited due to the large spacing between the primitives, leaving large pores in between. Thesepores are inter-connected under microscope, and are also surface-connected due to the aforementioned findings. Based on these observations, future effort will be made to reduce spacing between primitives (“Y Spacing” in ExOne R2 software) to improve the sintered density of copper.(a)(b)(c)Figure 11.

Based on these observations, future effort will be made to reduce spacing between primitives (“Y Spacing” in ExOne R2 software) to improve the sintered density of copper.(a)(b)(c)Figure 11. SEM of sintered AcuPowder 500A copper powder (1080 °C, 2 hours) at magnifications of (a) 200x, (b) 500x and (c) 2000x3.4 Effects of reduction on sinteringA reducing atmosphere not only improves the purity of final products by reducing copper oxide, but also has a positive effect on sintering and densification. The relative degree of surface oxidation affects the final sintered density of copper powders [32,33]. The proper reduction process for sintering copper powders is to fully reduce the oxidation and remove the water vapor before the closure of pores during sintering [34,35].

The proper reduction process for sintering copper powders is to fully reduce the oxidation and remove the water vapor before the closure of pores during sintering [34,35]. It has been found that an isothermal holdingbefore reaching the relative density of 85%, or a holding temperatures no higher than 900 °C for most powders is appropriate [32].To evaluate the effects of a reducing atmosphere, green parts made of AcuPowder 500A copper powder were sintered in controlled reducing atmosphere and non-reducing atmosphere (air with vacuum). The shrinkage and sintered density of these parts are compared in Table 5 . Compared to the non-controlled air atmosphere, the reducing atmosphere during sintering can improve the sintered density by more than 20% (Table 5). In the controlled reducing atmosphere, the densification was not significantly improved by increasing the concentration of hydrogen (Figure 12).

In the controlled reducing atmosphere, the densification was not significantly improved by increasing the concentration of hydrogen (Figure 12). While a reducing atmosphere can improve sintered density, the excessive supply of hydrogen than needed has not been demonstrated useful in improving densification.Table 5. A densification comparison between sintering AcuPowder 500A copper powder in reducing and non-reducing atmospheres (1060 °C, 2 hours)Volumetric shrinkage (%)Sintered density (% theoretical)Sintering in non-reducing atmosphere (air with vacuum) Sintering in reducing atmosphere (pure hydrogen)5.537.152.978.210050) l a c i t e r o e h t%(y t i s n e D d e r e t n i S95908580757065)%( e g a k n i r h S c i r t e m u o Vl48464442403836343260300%20%40%60%80% 100%0%20%40%60%80% 100%Hydrogen concentrationHydrogen concentrationFigure 12.

Sintered density and shrinkage of AcuPowder 500A copper powder in controlled reducing atmosphere (1080 °C, 2 hours, 100 ml/cc flow rate)3.4 Chemical compositionAs described in Section 2.2, chemical composition analysis was conducted via EDX spectroscopy. The results of the analysis of printed green parts (AcuPowder 153A copper powder) without any heat treatment and parts that were sintered in hydrogen are presented in Table 6. As observed in these results, the reducing atmosphere increased the purity of the finalcopper part, but was not able to completely convert copper oxide to its metal form during the sintering.Table 6.

Chemical composition comparison after sintering in a reducing atmosphereGreen parts Sintered partsAcuPowder 153A (1000 °C for 4 hrs.) 91.04% Cu 94.94% CuOzometal (1060 °C for 2 hrs.) 95.32% Cu 97.32% CuAcuPowder 500A (1080 °C for 2 hrs.) 96.36% Cu 97.11% Cu3.5 Uniaxial tensile testTo evaluate the strength of the printed copper parts, a tensile test specimen was first designedfollowing the ASTM E8 standard. The green part was printed larger to compensate for the shrinkage that would be caused by the sintering cycle (as determined in Section 3.2). The specimen was printed with AcuPowder 500A copper powder and was sintered at 1080 °C, 1060 °C and 1040 °C for 2 hours. As shown in Figure 12, the sintered specimen has a uniform shrinkage and the dimension of the sintered specimen was close to the ASTM standard.Figure 12. Tensile test specimen: sintered at 1060 °C (top) sintered at 1080 °C (bottom)Figure 13.

Tensile test specimen: sintered at 1060 °C (top) sintered at 1080 °C (bottom)Figure 13. Engineering stress-extension curve of copper P/M parts made via Binder Jetting in the tensile testingThe ultimate tensile strength of different specimen is summarized in Table 9. Compared to the engineering data of copper (ultimate strength of 210MPa and a yield strength of 70MPa), the copper parts made by Binder Jetting have up to 55.6 % of the theoretical ultimate tensile strength of copper. The specimen displayed certain ductility, as demonstrated by the curve. One can tell the tensile properties of copper P/M made by Binder Jetting is proportional to its sintered density. A reduced density has a tremendous negative effect on the tensile strength.Table 7.

One can tell the tensile properties of copper P/M made by Binder Jetting is proportional to its sintered density. A reduced density has a tremendous negative effect on the tensile strength.Table 7. Tensile stress of specimen sintered with various temperaturesSpecimen number 1 2 3 4Sintering profile 1080 °C, 2 hours 1060 °C, 2 hours 1040 °C, 2 hours 1040 °C, 2 hoursUltimate tensile strength (MPa) 116.7 73.3 8.4 6.9The relatively low strength can be attributed to the parts’ porosity, as previous studies have shown that the strength of porous sintered copper powder compacts is proportional to its porosity [2, 35]. To improve the mechanical strength of the copper fabricated via Binder Jetting, efforts need to be made in minimizing porosity or improving the densification during the sintering of copper.4. Summary and future workIn this paper the Additive Manufacturing of copper via Binder Jetting was explored.

Summary and future workIn this paper the Additive Manufacturing of copper via Binder Jetting was explored. Three purchased atomized copper powders were successfully printed on an ExOne R2 3D metal printer. The copper green parts were sintered in a tube furnace and a maximum of 85.5% of theoretical density has been achieved (Table 4). In the temperature range from 1060 °C to 1090°C the sintered density and shrinkage is proportional to the temperature (Figure 9). The reducing sintering atmosphere can achieve a purity of up to 97.3% (Table 6). The controlled sintering atmosphere with the presence of hydrogen can increase the purity of copper by up to 3.9% (Table 6) and can improve the sintered density by up to 25.3% (Table 5).

The controlled sintering atmosphere with the presence of hydrogen can increase the purity of copper by up to 3.9% (Table 6) and can improve the sintered density by up to 25.3% (Table 5). The tensile strength of printed sintered copper parts was 55.6% of the theoretical value due to the porosity (Table 7).In Binder Jetting, large particle size is preferred for powder spreading while fine powders are critical to the sinterability and have the potential advantages in improving resolution and surface quality. In was noticed one major challenge facing the Binder Jetting was to improve the sinterability and densification while the copper powders were relatively coarse and loosely packed. Therefore, future work will focus in the techniques that have the potential to improve copper sinterability while still maintaining good powder spreading. The authors will investigate multimodal mixture of powders, spray-dried granulated particles, and mixing sintering aids/inhibitors, etc.

The authors will investigate multimodal mixture of powders, spray-dried granulated particles, and mixing sintering aids/inhibitors, etc. In addition, the authors will explore the effect of printing parameters on sintered part density and characteristics. Further characterization will be focused in the areas of microstructural features, structural integrity, and thermal/electrical conductivity.5. AcknowledgementThis material is based upon work supported by the National Science Foundation under Grant No. #1254287. Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the views of the National Science Foundation. The authors acknowledge technical support provided by the ExOne Co. The authors also would like to thank Dr. Alex Aning and Dr. Tom Staley (Virginia Tech Department of Material Science and Engineering) for their assistance, and Ms.

The authors also would like to thank Dr. Alex Aning and Dr. Tom Staley (Virginia Tech Department of Material Science and Engineering) for their assistance, and Ms. Jennifer Sprouse – a researcher at Virginia Tech’s Research Experience for Teachers: Innovation-based Manufacturing program (NSF EEC #1200221) – for her assistance in conducting experiments.6. References[1] Bourell, D. L., Beaman, J. J., Leu, M. C., and Rosen, D. W., 2009, “A Brief History of Additive Manufacturing and the 2009 Roadmap for Additive Manufacturing : Looking Back and Looking Ahead.”[2] Badrinarayan, B., and Barlow, J. W., 1992, “Metal Parts From Selective Laser Sintering of Metal-Polymer Powders,” Solid Freeform Fabrication Symposium, pp. 141–146.[3] Badrinarayan, B., and Barlow, J. W., 1991, “Selective Laser Sintering of a Copper-PMMA System,” Solid Freeform Fabrication Symposium, pp. 245–250.[4] Murr, L. E., Gaytan, S. M., Ramirez, D. A., Martinez, E., Hernandez, J., Amato, K. N., Shindo, P.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperEXPLORATORY STUDY OF IN-SPACE WIRE ARC ADDITIVE MANUFACTURING WITH MODELING APPROACHWei Li*, Dan Bouzolin*, and Kishore Mysore Nagaraja*Department of Mechanical Engineering, The University of Texas at Dallas, Richardson, TX 75030AbstractCountries all over the world are rushing into space exploration due to crisis of energy and resources exhaustion on the Earth. Mars is an obvious target because it has a thin atmosphere, good geological similarity, and is close by in the Solar system. As the satellite of the Earth, Moon is another target since it is very close to the Earth. For the large spacecrafts such as Mars rovers, periodic maintenance is necessary to ensure the completion of long-duration exploration missions.

For the large spacecrafts such as Mars rovers, periodic maintenance is necessary to ensure the completion of long-duration exploration missions. In-space wire arc additive manufacturing (WAAM) provides a potential solution towards sustainable maintenance with onsite repair or additive manufacturing. For in-space manufacturing, reduced gravity is an important factor. In this work, WAAM processes under reduced gravity conditions on the Mars and Moon were studied through a multi-physics modeling approach. The metal droplet transfer, deposition geometry, thermal dissipation, and other key physics in WAAM were simulated.

The metal droplet transfer, deposition geometry, thermal dissipation, and other key physics in WAAM were simulated. To validate the modeling approach, an experimental case was conducted on an in-house WAAM platform under the Earth condition.Keywords: Wire arc additive manufacturing, thermal-fluid model, In-space manufacturing, reduced gravity, space exploration.IntroductionIn-Space Manufacturing (ISM) provides a solution toward establishing a sustainable human presence on Mars, the Moon, or space stations, through on-site additive manufacturing technologies (Fig 1). For ISM, reduced gravity is an important consideration in the manufacturing process. In this work, a Wire-Arc Additive Manufacturing (WAAM) technique was explored and validated against Earth gravity. A multi-physics model was used to simulate the WAAM process. The model was then subsequently applied to low-gravity environments such as on Mars and the Moon.

A multi-physics model was used to simulate the WAAM process. The model was then subsequently applied to low-gravity environments such as on Mars and the Moon. This work is an exploratory study for space exploration plans in the future.Fig 1 In-space manufacturing concept, with a close-up of an example welding processWire-Arc Additive Manufacturing (WAAM) is an advanced fabrication technique that uses an electrical arc to melt a metal wire for layer-by-layer deposition [1-5], as shown in Fig 2. WAAM can be used to create near1501net shape part geometries, making it highly desirable for material savings relative to subtractive processes like machining [6-8]. WAAM can also be used for repairing existing components as an alternative to traditional processes such as MIG and TIG welding [6, 9-11]. The WAAM process provides a high deposition rate, allowing it to produce parts quickly.

The WAAM process provides a high deposition rate, allowing it to produce parts quickly. The flexibility of WAAM in part geometry and rapid prototyping means that tangible cost and time savings can be achieved using this process[2, 9, 12-15].Fig 2 Overview of the wire-arc additive manufacturing process setupTo study the WAAM process under the reduced gravity conditions, modeling approach is a needed since it is challenge to conduct the WAAM experiment under the required reduced gravity environment. The WAAM process is normally difficult to model due to the complexity of the physical interactions involved, which include effects of electromagnetism, the Joule effect and radiation of the plasma[13, 14]. This research work aims to mitigate this fact by leaving out these complex electromagnetic effects and electrical arc modeling, instead approximating the heat source as a pulsed heat source with Gaussian energy distribution.

The benefit of this approach is that the overall simulation is less computationally expensive, allowing for rapid modeling and testing as well as larger part geometries.Methods, Models, and ExperimentThe model for this experiment consists of two major components: the metal wire (being fed from the robotic welding arm) and the substrate (onto which the metal wire is deposited). The wire and substrate material used for WAAM deposition is mentioned in Table 1. Both parts were modeled directly in Flow3D [16] as aluminum components. The model uses an electric heat source to simulate the heat input from the electrical arc of the welding machine. The fundamentals of the thermal-fluid, electric heat source, heat transfer, phase change, and other multiple physics were referred from the publication [14]. The Fig 3 as below shows the final part modeled in Flow3D with geometric parameters taken from the physical model.

The Fig 3 as below shows the final part modeled in Flow3D with geometric parameters taken from the physical model. The part geometry was modeled as a solid block of aluminum substrate (shown in blue) with a phantom region (shown in red) used to control the feed rate of the wire. The part was meshed in three regions: one center region (shown by the green lines) which has a cell size of 3e-4 m, and two side regions (shown by the cyan lines) which use a cell size of 5e-4 m. The higher fidelity in the center allows for more accurate modeling of surface tension while the coarser mesh on the outside keeps the simulation runtime down. Upon meshing, running, and rendering the results, the simulation setup functions as the experiment does with the wire feed, heating, and traversal.Fig 3 Flow3D model of final part geometry and mesh block delineations1502The physical parts were printed using the ABB IRB 140 robotic arm and Fronius TPS 320i electric welding system, as shown in Fig 4.

A wire diameter of 1 mm was used for the printing process. The wire feed rate is approximately 250 mm/s and the surface traversal rate is 30 mm/s. The power input into the system was calculated as 2154 W, which was obtained from the specifications used in the Fronius machine (136 A * 19.8 V * 80% = 2154 W) with 80% as the power efficiency. This power input was then pulsed at 20 Hz by using a transient power input file to mimic the behavior of the experimental setup. The substrate dimensions are 50 x 25 x 13 mm, with the wire traversal direction along the longest dimension of the substrate.Fig 4 Fronius welding system (left) and ABB robotic welding arm (right) used in the experimentThe experimental part dimensions and welding setup parameters were input into the Flow3D software to run the simulation.

The simulation was an iterative process, with parameters needed to be tuned in order to match the simulated model, but the final parameters lent themselves to relatively close correlation in simulated and experimental results. The material data of temperature-dependent properties of the Al 6061 alloy was used in the model. The simulation was run with Al 6061 wire, which has readily available data on how density, dynamic viscosity, thermal conductivity, and specific heat of the material all vary with temperature increase.Table 1 Summary table of parameters used in the simulationSubstrate Material Wire Material Wire Feed Rate Wire Traversal Speed Power Input Substrate Length Substrate Width Substrate Height Shield Gas Velocity Shield Gas DensityAl 6061 Al 6061 250 mm/s 30 mm/s 2693 W @ 20 Hz 50 mm 25 mm 13 mm 50 m/s 1.225 kg/m3The final parameters are listed in Table 1 as above, but several parameters had to be tuned in order to more closely match the experimental results.

For example, the wire feed rate and wire traversal speed have a significant effect on the overall behavior of the wire as it feeds down toward the substrate, as well as the physical material deposition process. With these parameters set too fast, the wire has no time to form droplets of molten metal, and instead feeds solid material directly onto the substrate. With the parameters set too slow, the metal droplet accumulates too much volume as it continues to be heated by the power source, and therefore does not continuously deposit, only depositing once the volume grows too large. Finally, the shield gas flowrate provides a downward force that pushes the droplet towards the substrate. With this parameter set too low, the droplet does not deposit onto the substrate, instead bubbling up on itself due to the force of the surface tension within the droplet.1503Results and DiscussionThe results of the Earth gravity simulation were first validated against experimental part geometry.

The experiment takes 1.7 seconds to deposit one layer of metal onto the substrate. The simulation also models 1.7 seconds of WAAM process, but takes about 20 hours to run. The main parameter of interest between the experiment and simulation is the overall layer geometry, which includes the layer width and height along the length of the substrate. The Fig 5 illustrates the experimental and simulated part geometries after undergoing the WAAM process.ZZYX13mmYFig 5 Experimental part geometry (left) vs simulated part geometry (right) using WAAM processThe experimental part geometry and simulated part geometry are in relatively close agreement, with a few minor differences. First, the experimental part geometry shown has an overall smaller bead width than the simulated part geometry, giving the appearance of a much larger aspect ratio in terms of the layer height.

First, the experimental part geometry shown has an overall smaller bead width than the simulated part geometry, giving the appearance of a much larger aspect ratio in terms of the layer height. In fact, the regions printed with the WAAM process are similar in size between both the simulation and experiment. The Fig 6 as below shows a side view with a more comparable look at the printed part geometries.ZZXXFig 6 Side views comparison of experimental part geometry (left) vs simulated part geometry (right)In addition, the front view of the part geometries is shown in Fig 7 as below, illustrating the width of theprinted features resulting from the WAAM process.2.473 mm1.752 mm2.625 mm3.942 mmFig 7 Front view comparison of experimental part geometry (left) vs simulated part geometry (right)From an initial inspection, the experimental part geometry appears slightly taller while the simulated part geometry appears slightly wider.

To quantify these relative differences between the simulated and experimental1504Xparts, the sizes of the printed features were measured. Table 2 below shows the quantitative difference in printed part geometry between the experimental and simulated parts.Table 2 Summary table of printed part geometrySample Type Experimental SimulatedMaximum Print Height (mm) Maximum Print Width (mm) 2.42 2.093.70 4.78To understand these differences, it is important to understand how the WAAM process works: more specifically, how material is deposited onto the substrate. The primary mechanism by which metal is melted and deposited onto the substrate is through a droplet-based deposition process. The heat source melts the incoming wire, which bubbles up into a droplet with nearby liquid metal.

The heat source melts the incoming wire, which bubbles up into a droplet with nearby liquid metal. Once the droplet is big enough, it detaches from the wire and deposits onto the substrate below.(a)(b)(c)Fig 8 Droplet-based deposition process with (a) initial wire, (b) material accumulation, and (c) droplet depositionAs mentioned before, the experimental part geometry had slightly taller ridges while the simulated part geometry had slightly wider ridges. The reason for this difference becomes apparent upon closer inspection of Fig 8, which depicts the droplet-based deposition process. Though this minor difference exists between the experiment and simulation, the overall modeling of the WAAM process and droplet deposition were sufficient to continue with further simulation. Once the results of the Earth-gravity simulation were validated against the experimental part geometry, a baseline was established for which to compare future simulations to.

Once the results of the Earth-gravity simulation were validated against the experimental part geometry, a baseline was established for which to compare future simulations to. To simulate the physics of acceleration due to gravity, the z-component of gravity in the simulation was set to the various values ranging from Earth gravity (-9.81 m/s2) to Moon gravity (-1.62 m/s2). A summary of the parameters that were updated for this investigation is shown in Table 3 below.Table 3 Summary table of gravity parameters used for different simulationsSimulation Number 1 2 3 Upon changing the gravity, the simulation was re-run for each scenario.

The results for both the MarsSimulated Environment Earth Moon MarsAcceleration due to Gravity (m/s2) -9.81 -1.62 -3.72gravity simulation and Moon gravity simulation are summarized in the Fig 9 as below.1505Fig 9 Comparison of Mars gravity simulated result (left) and Moon gravity simulated result (right)Upon inspection, the two results show close correlation with each other as well as with the Earth gravity result. The overall part geometry remains similar between trials, indicating that the change in gravity may not be the driving factor in droplet deposition. From the droplet-based deposition process in Fig 8, the droplet deposits onto the substrate upon growing large enough to touch the surface, at which point the contact force from touching the substrate overcomes the surface tension within the droplet keeping it together.

Because the droplet does not break off prior to touching the surface, the force of gravity does not have as significant of an effect as other forces such as the wire feed rate and downward force due to the shield gas.Fig 10 Front view of Mars simulation result (left) and Moon simulation result (right)The front views of each simulation (Fig 10) show close correlation to one another as well as the Earth gravity result. Though the gravity varies from Earth gravity by nearly 3x for the Mars simulation and 6x for the Moon simulation, the overall part width remains consistent.

Though the gravity varies from Earth gravity by nearly 3x for the Mars simulation and 6x for the Moon simulation, the overall part width remains consistent. The heights of the resultant printed features remain consistent between simulations, which can be validated further using the side views of each part.Fig 11 Side view of Mars simulation result (left) and Moon simulation result (right)From the side views of both low-gravity simulations (Fig 11), it becomes clear that the droplet formation process develops features with similar heights, with only slight variation between each ridge as to where the height grows smaller or larger. The side view also provides valuable insight into the location at which droplets1506are formed. If gravity were to add a substantial force pulling down the droplets toward the substrate, it would be expected to see droplets forming smaller ridges but more frequently along the substrate.

If gravity were to add a substantial force pulling down the droplets toward the substrate, it would be expected to see droplets forming smaller ridges but more frequently along the substrate. However, the droplets form in the same locations each time, giving rise to consistently manufactured ridges between each case. We can further inspect the top view to validate the droplet formation locations along the substrate.Fig 12 Top view of Mars simulation result (left) and Moon simulation result (right)Comparing the top views of both low-gravity simulations (Fig 12), it shows that the droplets formed during the WAAM process detach at approximately the same locations and at the same time in both simulations.

The results indicate that the WAAM process has no issue with droplet formation in such low-gravity environments, owing to the fact of several forces acting in tandem on the droplet in addition to gravitational force.ConclusionThe results of the Earth gravity simulation were first validated against experimental part geometry. The electric heat source was effective at heating the wire for melting, as evidenced by the general wire melt pattern matching between the experimental and simulated results. However, improvements could be made to approximate the wire-arc behavior, as the pulsed electric heat source does not transfer as much heat to the substrate in the simulation compared to the wire arc in the experiment.As evidenced from the figures above, varying the acceleration due to gravity did not have a significant effect on the simulation results.

This suggests that the droplet-based deposition process may not be driven strictly due to gravity, but may be driven by other, more significant factors, such as the wire feed rate and the force on the fluid due to the incoming shield gas.The simulated results show that the wire-arc additive manufacturing process may be extended to in-space manufacturing with few differences from what is seen currently on Earth. However, caution should be exercised in extending this result, as the simulation and experiment only correlated the first layer of the WAAM process. More work can be done to identify how the parts vary with multiple printed layers.In general, this study provides an overview of the wire-arc additive manufacturing process and the effects of varying gravity on this process.

Preliminary work shows that the wire-arc additive manufacturing process is suitable for in-space manufacturing, with only minor differences in part geometry between parts printed in Earth gravity versus Mars or Moon gravity. Future studies can explore changing the heat source to approximate the wire-arc more closely in the simulation as well as printing multiple layers at a time, both of which would provide valuable information about the validity of WAAM for in-space manufacturing.1507Reference[1] C. Cunningham, J. Flynn, A. Shokrani, V. Dhokia, S. Newman, Invited review article: Strategies and processes for high quality wire arc additive manufacturing, Additive Manufacturing 22 (2018) 672-686. [2] S.K. Dash, Study of Wire Arc Additive Manufacturing With Aluminum Alloy 2219, (2019). [3] Y. Ali, P. Henckell, J. Hildebrand, J. Reimann, J. Bergmann, S.

Li, A multi-bead overlapping model for robotic wire and arc additive manufacturing (WAAM), Robotics and Computer-Integrated Manufacturing 31 (2015) 101-110. [14] S. Cadiou, M. Courtois, M. Carin, W. Berckmans, 3D heat transfer, fluid flow and electromagnetic model for cold metal transfer wire arc additive manufacturing (Cmt-Waam), Additive Manufacturing 36 (2020) 101541. [15] D. Ding, Z. Pan, D. Cuiuri, H. Li, S. van Duin, N. Larkin, Bead modelling and implementation of adaptive MAT path in wire and arc additive manufacturing, Robotics and Computer-Integrated Manufacturing 39 (2016) 32-42. [16] W. Li, M. Kishore, R. Zhang, N. Bian, H. Lu, Y. Li, D. Qian, X. Zhang, Comprehensive studies of SS316L/IN718 functionally gradient material fabricated with directed energy deposition: Multi-physics & multi-materials modelling and experimental validation, Additive Manufacturing 61 (2023) 103358.1508.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEXPLORING A SUPERVISORY CONTROL SYSTEM USING ROS2 AND IOT SENSORS M. A. Roach1, J. Penney1, B. H. Jared11Mechanical, Aerospace, and Biomedical Engineering, University of Tennessee, Knoxville, TN 37996AbstractWhether collecting data from process monitoring sensors or controlling a system of multiple actuators and electrical systems, a powerful supervisory control system must be developed for additive manufacturing (AM) systems. The Robot Operating System version 2 (ROS2) is a set of software libraries that can be used to control robotics systems and has tools for sensor value publishing.

The Robot Operating System version 2 (ROS2) is a set of software libraries that can be used to control robotics systems and has tools for sensor value publishing. This research project is exploring the use of computational nodes connected to process monitoring sensors and robotic or electrical systems to allow for a more in-depth knowledge of the system health and process as well as open the possibilities of process control. These nodes can be connected and controlled by the ROS2 architecture. Work will be discussed exploring the reliability and speed of common AM processes and sensors such as robot controllers and thermal monitoring.BackgroundWhen switching from traditionally cast metals to AM created metals, the properties of the metal are no longer guaranteed and may not be homogenous throughout the volume of the material.

Due to this downside in AM, the need for process monitoring is now formed, as a method of watching signals and changes during the metal deposition process can be beneficial and may indicate process conditions have deviated significantly enough to allow a defect to form in the material.In this research, the AM methods of interest are wire-fed directed energy deposition (DED) such as wire-arc AM. DED is traditionally done by programming 2- and 3-dimensional paths which follow sequential slices of a pre-defined geometry on top of a 2- and 3-dimentional substrate. These paths are then used by robots or CNC equipment to manipulate a deposition head along the path or manipulating the build area about the deposition head along the pre-defined path.

These paths are then used by robots or CNC equipment to manipulate a deposition head along the path or manipulating the build area about the deposition head along the pre-defined path. The deposition head consists of an energy source to melt the material with a wire-feeding mechanism to supply the melt pool with new material for the purpose of deposition.While this paper does not focus on a specific DED setup or system, the methods and results are applicable not only to many DED systems but also many other AM systems in general. The reason for this is that many AM systems rely on similar sensors to monitor the deposition process. Such sensors can include thermal monitoring, robot or CNC equipment programmed and actual position monitoring, 2- or 3-dimensional imaging sensors, and power monitoring.

Such sensors can include thermal monitoring, robot or CNC equipment programmed and actual position monitoring, 2- or 3-dimensional imaging sensors, and power monitoring. There exist many methods for AM system controls, collection of sensor information, open- and closed-loop control, as well as path planning to allow the workflow from pre-defined geometry to final part. In this paper, the specific system under investigation is defined in the remainder of this section.This specific system consists of a KUKA KR6-2 robotic manipulator with a Fronius TPS 4000 CMTAdvanced robotic weld head. This welding head uses the pre-defined welding profiles generated by Fronius and are executed by the KUKA ArcTech package in job mode. The robot path is generated using Octopuz path planning and robot simulation software where the final files (.src and .dat) are copied over to the KUKA controller and run by the operator to print the desired geometry.

During the printing process, a manual pause is added between each sequential layer to prevent the heat in the part from building up. This is normally done in one of two ways; either a 120 second delay is manually coded into the path plan using Octopuz, or the program is manually paused by the operator after each layer to allow for a manual temperature measurement using a handheld non-contact IR thermometer to monitor the temperature of the deposited material until it passes a pre-1094defined threshold. This threshold varies for different materials and is not standardized. The following sections will explore the use of ROS2 and IoT sensors to work towards the automated thermal monitoring and path plan resuming without the need for operator intervention.This system is a development system; therefore, it integrates many different sensors and process monitoring equipment.

This suite of sensors is always changing based on research needs, but for this research the configuration is as follows. The robot is monitored using the KUKA.RobotSensorInterface (RSI)[1] package which reports back current tool positions as well as error dialogue and robot status. This is connected to the existing Windows 10-based control computer via an ad-hoc ethernet network. Also connected to this network are two cameras, a FLIR A50 thermal camera and a XIRIS XIR-1800 welding camera. These cameras watch the bulk temperature distribution across the deposited material and the arc and melt pool dynamics and temperature, respectively. Finally, the remainder of the sensors are currently integrated with a custom LabVIEW VI that includes 4 k-type thermocouples, welding voltage and current, and the previously mentioned RSI.

Finally, the remainder of the sensors are currently integrated with a custom LabVIEW VI that includes 4 k-type thermocouples, welding voltage and current, and the previously mentioned RSI. The two cameras listed above are controlled by their own software and do not integrate with the LabVIEW system.Issues have occurred frequently with the maintenance of the LabVIEW controller VI that include driver mismatches, the lack of timestamp recording for time-series measurements, and the ability to easily add such a feature into the existing VI. Due to the way in which the VI is executed, the sensor acquisition and control is collected on a time-defined loop and lacks the ability for a more intelligent data collection method. Along with these issues, the need to purchase LabVIEW licenses and National Instruments equipment quickly increases the cost of development of these systems.

Along with these issues, the need to purchase LabVIEW licenses and National Instruments equipment quickly increases the cost of development of these systems. Especially when a specific sensor, such as the welding camera, does not have existing compatibility with LabVIEW. ROS2 will be explored to determine if it can solve these problems and still deliver adequate performance.ROS2 IntroductionAs their website states “ROS (Robot Operating System) is an open-source software development kit forrobotics applications. ROS offers a standard software platform to developers across industries that will carry them from research and prototyping all the way through to deployment and production.”[2] This software has been used in many AM systems at many institutions and has proven to be a valuable tool in this research space. The ROS ecosystem advertises 8 main benefits on their website and some of these benefits were seen directly with this research.

The ROS ecosystem advertises 8 main benefits on their website and some of these benefits were seen directly with this research. These benefits include a substantially large and active global community of developers, a continued and proven use in systems spanning single-student projects to large-scale collaborations and competitions, customizable and shorter time to market, multi-domain due to the generality of the software model, multiplatform so it runs on systems from Mac and PC to Linux and embedded platforms and numerous programming languages from C++ and Python to MATLAB, completely open source, non-restrictive Apache 2.0 license, and finally industrial support specifically from ROS2[3]. While not all these claims can be verified in this research due to the limited scope and timeline, future work as well as existing ROS-based projects continue to prove its viability.ROS2 comes in two major versions, ROS and ROS2.

ROS2 is the newest version and focuses on theindustrialization of the software framework. ROS and ROS2 operate individually but can still communicate via the ROS1 to ROS2 bridge. This is especially helpful when existing packages and libraries have not yet been migrated by their developers. This research will focus entirely on ROS2.The ROS2 environment is composed of discrete elements, all with a defined task such as data acquisition for one sensor or a decision-making algorithm output. To begin, all work is done within a Workspace[4]. This is a simple folder with defined subfolders like src for organizing your system as a whole. For example, a complete AM system would operate within one workspace if only one computer was used. The next organizational element is a Package[5]; these are the backbone of ROS2 and its modularity.

For example, a complete AM system would operate within one workspace if only one computer was used. The next organizational element is a Package[5]; these are the backbone of ROS2 and its modularity. Packages are collections of executable scripts and resources to control a specific entity like a sensor, robot, or specific system, for example existing packages can be used to interface with the KUKA KR6-2 and RSI. Within these1095organizational elements and executable scripts, the main ROS2 elements are called and run to operate the target system.The most important element is a Node[6], these are responsible for a singular and modular purpose, forexample collecting data from a sensor, or conducting computations or machine learning classifications. Usually, each Node is initialized by a single executable file, like a Python script or compiled C++ executable, using object-oriented programming. Similarly multiple Nodes can be initialized in a single executable file.

Similarly multiple Nodes can be initialized in a single executable file. The next element is the Topic[7]. These are how data is moved between Nodes. These are defined within the Nodes to publish and subscribe to specific Topics. A simple example is shown below in Figure 1, where three nodes are running and publishing or subscribing to one Topic by sending and receiving the Message. A fourth element is shown called “ros2 topic echo” which is an example of using the command line interface ability of ROS to print the published Topic Messages to the terminal.Figure 1: Basic ROS2 structure[8]More advanced capabilities exist in ROS2 that expand the functionality to cover nearly all actions; this is done using Services[9] and Actions[10]. Services provide data when they are specifically called by a Node. These can be used for actions like checking an enabling signal before an operation is executed, but the enabling signal does not need to be published continually to rest of the system.

These can be used for actions like checking an enabling signal before an operation is executed, but the enabling signal does not need to be published continually to rest of the system. Finally, an Action is a system task with a goal, feedback, and a result. Using an Action is helpful when the system must do something such as move the tool to a pre-defined location; this action will send the target coordinates, return the coordinates as the tool is moved, then the final status of the system is sent upon completion of the goal. An example of this system can be seen in Figure 2, where the same nodes from Figure 1 are used, but with different elements. Note, that all Nodes, Topics, Services, and Actions in this example can coexist in one environment, rather than how they are depicted as two systems.1096Figure 2: More advanced ROS2 system[8]There exist other elements and capabilities within ROS2, but in this research only Nodes and Topics willbe used.

The authors acknowledge that for a complete AM system, all elements must be used for the most optimal supervisory control system.An additional comment can be made in regards to the use of generative AI, such as ChatGPT[11], forreducing the time required to develop ROS2 nodes. For example, the following prompt can be used to generate a starting point for a Python script that initializes a node, connects to a specific sensor, and publishes that sensor value to a topic: “write a python script for a ROS2 node that connects to ‘specific sensor’ and publishes the value as a 32-bit float message”. If the specific sensor has a public development kit that has a python library option, and the generative AI has been able to learn about it, the response should allow for a more streamlined start to a node script.

This may not work in all cases and may also require further program troubleshooting and installation of required libraries.MethodsTo test the capabilities of the ROS2 software, an experimental setup is created. This consists of a simplesystem designed to simulate an AM system thermal monitoring setup and can be seen in Figure 3. The major components are a Raspberry Pi 4b 8GB (RPi), 500C Ohaus hotplate, and an Omega OS37-10-K-HIE adjustable IR sensor with k-type thermocouple output. By heating a target material on the hotplate to a representative temperature of 90C, the RPi can run the ROS2 system to test the basic functionality by interfacing with the pyrometer for simple temperature monitoring.1097Figure 3: Simple testbed systemThe RPi is connected to a standard breadboard using a ribbon cable and a general-purpose input/output(GPIO) cobbler. This allows for the interface pins on the RPi to be made available on the breadboard rows.

This allows for the interface pins on the RPi to be made available on the breadboard rows. This system must be able to interface with the pyrometer, collect time series data, and display the pyrometer temperature to a theoretical operator. To interface with the pyrometer an Adafruit MAX31855 thermocouple amplifier breakout board is connected to the breadboard. This breakout board communicates with the RPi via Serial Peripheral Interface (SPI) by using the 3.3 volt and ground pin for power, SCLK and MISO pins for SPI interface, and GPIO 5 for data. To display the current temperature, a basic 2x16 LCD screen is connected to a basic Inter-Integrated Circuit (I2C) board. This board connects to the RPi using the 5 volt and ground pins, and the SDA/SCL data and clock pins. Figure 4 shows a full wiring diagram of the RPi and its peripherals.1098Figure 4: Electrical connections for testbed systemROS2 ArchitectureWith hardware connected property, the ROS2 software can now be developed.

The RPi is flashed with the Ubuntu Desktop 22.04 LTS 64-bit image[12]. The operating system is set up using standard configuration, Secure Shell Protocol (SSH) enabled, and networked using Wi-Fi. SPI and I2C interfaces are enabled through raspi-config. The RPi is accessed using Microsoft Visual Studio Code[13] via SSH for the remainder of development. In order to execute the node responsible for communicating with the pyrometer over SPI, the Ubuntu user invoking the ROS node needs access to the “/dev/spidev0.0” access group, found by using the terminal command “ls -l /dev/spidev0.0”.ROS2 Humble is installed using the standard Ubuntu Debian package “ros-humble-desktop”[14]. A testbed workspace was created in the home drive of the main user. Now, the main structural elements of the ROS2 environment can be created.

A testbed workspace was created in the home drive of the main user. Now, the main structural elements of the ROS2 environment can be created. Four packages were created for this testbed and include: Pyrometer Package to interface with the pyrometer and breakout board, LCD Package to display values on the LCD screen, HDF5 Package to archive the temperature time series data as an HDF5 file, and a Bring-Up Package to make it easier to start the testbed system. This general environment can be seen in Figure 5.1099Figure 5: Testbed software structure and data flow diagramEach package is modular and can be used by any system that uses the same sensor or interface. Forexample, all AM systems that use the pyrometer (or, more realistically, simply the MAX31855 breakout board) can use the Pyrometer Package.

Forexample, all AM systems that use the pyrometer (or, more realistically, simply the MAX31855 breakout board) can use the Pyrometer Package. The only exception is the Bring-Up Package which is specific to the system configuration and contains the ROS Launch file, which can be thought of as the orchestrator of the system, that selects and runs the required nodes to boot-up the system as expected. These packages are all simply folders within the “src” folder of the Workspace and themselves contain a “src” folder that contains the executable files that run each Node of the system. These package folders are synced with GitHub[15] to maintain versioning and monitor changes.The data flow diagram seen on the right of Figure 5 illustrates how the topic data is sent between the nodes and then how that data is collected and displayed through system inputs (pyrometer value) and outputs (HDF5 data package and LCD screen).

Once the Launch file is executed, all nodes and topics will function to run the system as defined. This testbed system uses Python 3.10 to program the nodes for all packages. Each node contains a single executable Python script as seen in Figure 6.Figure 6: Python files and their contentsWithin the Pyrometer Package, the python file (“max31856_read_node.py”) utilizes the Adafruit MAX31856 Python library, specifically written for this breakout board, and installed using the command “sudo pip3 install adafruit-circuitpython-max31855”[16], [17]. This library connects to the board and then creates a Python class object that can be called to acquire a temperature measurement. Additional functionality includes the definition of the pyrometer minimum and maximum measurement limits as well as error handling for communication and thermal bounds of the system. The Python script also initiates a topic publisher with the topic name “temperature” and a message format of a 32-bit float number.

The Python script also initiates a topic publisher with the topic name “temperature” and a message format of a 32-bit float number. The MAX31855 python library uses a1100delay in the acquisition of a temperature so this node has the temperature measurements collected and published as fast as the library allows.Within the LCD Package, the python file (“lcd_node.py”) utilizes the I2C library[18] and a library forcontrolling the LCD via the 16-pin interface[19]. The script also initializes a topic subscriber, subscribing to the topic of “temperature”. Every time the subscriber is given a new message from the topic, it runs a callback function that updates the LCD display with the new temperature, using the LCD and I2C libraries.Within the HDF5 Package, the python file (“data_archiver_node.py”) and utilizes the HDF5 library[20].This script initializes an HDF5 file with the correct structure and metadata.

Similar to the LCD Package, the script initiates a topic subscriber that runs a callback function that adds the temperature value and timestamp to the HDF5 file. Once the node is terminated, the HDF5 is closed and ready for retrieval.Within the Bring-Up Package, the python file (“testbed_bringup.launch.py”). As mentioned previously,this package is custom for each system. In this script, the launch sequence is relatively simple as it only launches the three main nodes (pyrometer, lcd, and data archiver). These nodes can be run using the terminal and normal ROS2 interface commands, but using the launch file reduces the need to open three terminals and run each node separately.Experimental MethodsThe final experiment for this research was to record data while completing the pyrometer calibrationprocedure. First, the RPi was connected to the breakout boards and pyrometer, then powered on.

First, the RPi was connected to the breakout boards and pyrometer, then powered on. The hotplate was loaded with a sample material of steel with steel weld beads and the pyrometer was positioned directly above and set at the correct height for the pyrometer measuring distance and set to read 75% of the target temperature of 90C. The ROS2 launch file was then started. The hotplate setpoint was adjusted to 90C and allowed to reach setpoint and remain stable for many minutes. The pyrometer was then calibrated to show the correct reading of 90C then the hotplate was turned off and allowed to cool to ambient temperature over many minutes. The ROS2 network was then terminated and the HDF5 file was recovered.ResultsThe resulting time series data for temperature from the HDF5 file is shown in Figure 7. The initialtemperature can be seen raising from a starting point of about 55C and leveling out at slightly less than 70C.

The initialtemperature can be seen raising from a starting point of about 55C and leveling out at slightly less than 70C. After the spike in temperature caused by the manual calibration procedure, the temperature reads 90C. As the hotplate is cooled, the pyrometer temperature settles to slightly above 70C, indicating either further calibration processes are needed, or the thermocouple amplifier is not operating as expected.Final temperature data was collected at around 4Hz which is most likely limited by the Adafruit Python library using a hard-coded delay. A higher data collection frequency may be possible by changing developing a more efficient Python function that collects data without a defined delay.

A higher data collection frequency may be possible by changing developing a more efficient Python function that collects data without a defined delay. This data collection frequency is high enough for many AM sensor needs, but other sensors will need higher throughputs for both speed and message size.1101Figure 7: HFD5 file interface and temperature graphThe purpose of this research was to test the feasibility of using ROS2 as a process monitoring and control system for AM. This research shows the setup of a ROS2 environment to be relatively simple and well documented in support and online communities. Rather than writing Python scripts for both sensor data acquisition and system operation, only the sensor interface needs to be written, while ROS2 manages operation.

Rather than writing Python scripts for both sensor data acquisition and system operation, only the sensor interface needs to be written, while ROS2 manages operation. Additionally, the Python scripts defining ROS2 nodes can be significantly short and quick to write, as initializing a ROS node takes only a few lines of code and with an existing Python library sensor integration may only take one or two more lines of code. Given the highly modular design, packages can be easily interchanged throughout different AM systems, and unique versions of nodes can be run from the same package, as defined by the needs of the system. Such new node version can also be made available to other systems that use that package, thus expanding the features of all systems.ConclusionThis research has shown that the collection and recording of pyrometer sensor data at 4Hz is possibleusing ROS2 and allows for highly modular and scalable systems of increasing complexity.

Future research will focus on the development of additional packages for sensors such as the FLIR A50 thermal camera and XIRIS XIR-1800 which have existing software development kits (SDKs) which allow for integration into the ROS environment. Further research also needs to be done for developing more functionality through decision-making nodes, possibly powered by machine learning. Work is also being done to implement HDF5 files and define a more standardized AM data management framework. Finally, the functionality of ROS2 systems can be expanded by utilizing the ability to network multiple computers running ROS. Bringing data computation closer to the source of creation will allow for a more efficient system that is not controlled by a centralized computer running all actions, acting as a single point of failure.AcknowledgementsThe figures and text in this report are human generated. The python scripts used in this research werepartially developed using generative AI[11].

The research aimed to produce dense parts using both scanning strategies, employing custom Python code for raster melt beam path generation and PixelMelt software for stochastic spot melting path generation. After optimizing process parameters, 10mm height builds for each scanning strategy were fabricated, and their microstructure, hardness, and density were analyzed using optical microscopy and SEM, Vickers microhardness scale, and a pycnometer.

The findings reveal valuable insights into the effects of scanning strategies on the microstructure, hardness, and density of IN718 alloy components, advancing additive manufacturing knowledge.NomenclaturePBF – Powder Bed Fusion EB-PBF – Electron Beam Powder Bed Fusion IN718 – Inconel 718 (Nickel Based Alloy) µm – Microns µs - Microseconds mm - Millimeters Ea – Energy Per Area (J/mm2) °C – Degree Celsius Eas - Energy Per Area Spot(J/mm2)1 IntroductionThe aerospace and defense sectors are in constant pursuit of materials that exhibit excellent strength and resistance to deformation at high temperatures. Nickel-based superalloys, particularly Inconel 718 (IN718), have gained considerable attention due to their superior strength, versatility, and resistance to corrosion even at elevated temperatures up to 650°C [1]. Additionally, these alloys display commendable weldability and formability.

Additionally, these alloys display commendable weldability and formability. This alloy is mostly composed of nickel, iron, chromium, niobium, molybdenum, titanium, cobalt, and small amounts of aluminum and iron. The exact composition ranges from 50-55% nickel, 17-21% chromium, 4.8-5.5% niobium, 2.8-3% molybdenum, 0.65-1.15% titanium, 1% cobalt, with the remainder being iron and aluminum [2,3]. Its phase composition primarily includes a γ matrix, enriched with precipitates of γ', γ", δ, and certain carbides [4,5].IN718 alloys are widely in focus due the the capability of them being produced additively through metal Additive Manufacturing. The most common and widely used metal AM technologies are – Powder Bed Fusion (PBF), Directed Energy Deposition (DED), and BinderJetting (BJT). Among them, PBF has already wide459recognition for producing IN718 parts. EB-PBF is one type of PBF machine that uses an electron beam in a vacuum chamber and melt the materials layer by layer.

Among them, PBF has already wide459recognition for producing IN718 parts. EB-PBF is one type of PBF machine that uses an electron beam in a vacuum chamber and melt the materials layer by layer. In recent years, a sum amount of research has been conducted using EB-PBF to produce IN718 parts using commercial EB-PBF machines exploring novel scanning strategies to control the microstructure [6,7,8,9].This research work will evaluate the fabrication of IN-718 alloy using an open source EB-PBF machine called FreemeltOne (Freemelt AB, Sweden) using two different scanning strategies : Bi-directional raster and stochastic spot melting. The goal of this study was to identify differences in terms of microstructure, hardness and density for IN718 in raster scan and spot melt scanning strategy.2 MethodologyFeedstock CharacteristicsThe powder utilized for this research was plasma-atomized, spherical IN718 powder obtained from vendor AP&C.

It exhibited a particle size distribution ranging from 45 to 106µm, with respective D10, D50, and D90 values measured at 52µm, 73µm, and 105µm. The powder's flowability was measured using a hall flowmeter following the ASTM B213-20 standard and was found to be 11s/50g, indicating satisfactory flowability for the PBF application.EB-PBF SystemThe machine used for part fabrication was FreemeltOne EB-PBF machine (Freemelt AB, Sweden). It uses CO2 laser to heat up the cathode (LaB6) in vacuum and uses electron beam to melt metal powders layer by layer. The machine operates at 60kV in a vacuum with base pressure ranging from from10-6 hPa (mbar) to 10-7 hPa (mbar). The maximum beam power for this machine is 6kW. The print process for the FreemeltOne is divided into four stages, namely startheat, preheat, melt and postheat, which are executed sequentially in each cycle during printing.Process ParametersStainless steel build substrate was used as build plate.

At the beginning of the building, the plate was heated to 1050°C using a custom-coded beam pattern. After that, as powders were spread of 0.75 µm layer thickness to the build plate, a sequence of preheat patterns was employed to preheat the powder to minimize smoke events. Next, melt was done and the process repeated itself.For the Melt process, two different scanning strategies: Bi-directional Vector Raster (Figure 2.1a) and Stochastic Spot Melt (Figure 2.1b) were implemented.(a) Bi-directional Raster ScanN and n+1 layers(b) Stochastic Spot melt n and n+1 layersFigure 2.1 Schematic diagram for Bi-directional Raster and Spot Melt Strategy460For the raster scan – python script was written to generate open beam path file (.obp) which dictates the beam path. For the Spot melt, cloud-based PixelMelt software (Freemelt, AB, Sweden) was used to prepare the build obp file for the melt. In both cases, a build of 10mm height was done.

For the Spot melt, cloud-based PixelMelt software (Freemelt, AB, Sweden) was used to prepare the build obp file for the melt. In both cases, a build of 10mm height was done. For raster, the shape of the build was square of 15x15mm and for spot melt it was circular geometry of diameter 15mm.For the raster and the spotmelt parameters for melt is explained by Table 2.1 below:Table 2.1: Process ParametersScanning StrategyRaster MeltStochastic Spot MeltParameter Type Power Speed Line Offset Ea (J/mm2) Power Mesh Size Beam Dwell Eas (J/mm2)Values 300W 800mm/sec 0.1mm 3.75J/mm2 360W 0.1mm 0.250 milliseconds 9J/mm2Microstructure CharacterizationTo analyze the microstructure of the builds, sectioned pieces were cut along the build directions and then hot-mounted using an ATM Opal 460 (Haan, Germany). These samples were further prepared by mounting and polishing them with the help of an ATM SAPHIR 530 (ATM GmbH, Germany) semi-automatic system.

These samples were further prepared by mounting and polishing them with the help of an ATM SAPHIR 530 (ATM GmbH, Germany) semi-automatic system. To reveal the microstructure, Kralling's No. 2 etchant was used. The etched samples were then observed using a Keyence VHX 7000 (Keyence Corp., Japan) Series digital microscope. To investigate deeper into the details of grain boundaries, Scanning Electron Microscope (SEM) images were taken using a Jeol SEM machine, with the imaging plane oriented in the build direction.Density and HardnessThe density of the built samples was measured using the Accupyc II 1340 (Micromeritics Instruments, USA) series pycnometer, which utilizes helium as its operating gas.

For the hardness measurements, the CHD Master microhardness tester (QATM, Germany) with the Vickers scale (HV1) was employed.3 Result and DiscussionMicrostructure CharacterizationThe as-polished microstructures, oriented along the build direction, for both the raster and stochastic spot melt techniques, are illustrated in Figure 3.1. Raster melt build had lower porosity than stochastic spot melt- observed in as-polished microstructure.461YZa) Raster Melt (contrast and sharpness enhanced)XZYZb) Stochastic Spot Melt (contrast and sharpness enhanced)Figure 3.1 (a) As-Polished (Raster) ; (b) As-Polished (Stochastic Spot Melt)During the stochastic spot melt process, higher energy was required to obtain a dense build which might contributed in resulted porosity.

It was observed that – if the same Eas was maintained, by altering high power, lower spot dwell time, or low power and higher spot dwell time – two different cases occurred.For case I, lower power higher spot dwell, a concentrated porosity was at the center of the specimen in the build direction (Figure 3.2a). For case II, high power and lower spot dwell time, small porous regions were observed all over the specimen (Figure 3.2b).

However, stitched images of stochastic spot melt in the build direction (Figure 3.3b) provide insights that in the edges, the grains are less columnar than the center. This could be the faster cooling rate of the edges of the specimen than the center, which yielded fewer columnar microstructures than the center.463Further insights were found in grain microstructures at the edges (Figure 3.4) for stochastic spot melt builds, when compared in case I and case II conditions as elaborated earlier. These images were taken in optical shadow mode keeping the light angle in the same direction of the etched samples. It is observed that, at near edges, even though the Eas was kept the same during the build, lower dwell time resulted in fewer columnar grains at the edges than the higher dwell time builds.

It is observed that, at near edges, even though the Eas was kept the same during the build, lower dwell time resulted in fewer columnar grains at the edges than the higher dwell time builds. This provides us a context that it might be possible to refine microstructures by properly altering dwell time with power in case of spot melt.XZXZa) 540W |170μs Dwell | 9.18J/mm2 Easb) 300W |300μs Dwell | 9J/mm2 EasXZc) 360W |250μs Dwell | 9J/mm2 EasFigure 3.4 Etched images near edges for 3 different Process parameters while maintaining similar EasAdditional investigation through SEM was done to gain insights of the the grains for both strategies using backscattered (BSE) and secondary electron (SE) mode (Figure 3.5). Even though, BSE and SE mode were tweaked – not enough information for grains were found rather except the confirmation of columnar growth of grains in both cases.

Even though, BSE and SE mode were tweaked – not enough information for grains were found rather except the confirmation of columnar growth of grains in both cases. However, stochastic had more less visible columnar grain than raster melt.Figure 3.5 SEM images in SE and BSE Mode to evaluate microstructure Raster (Left) ; Stochastic Spot Melt (Right) -contrast, sharpness and brightness boosted464Density and Hardness The raster build achieved a density of 99.71% according to pycnometer measurements, and ImageJ analysis indicated a range of 99.87-99.91%. Meanwhile, for the stochastic spot melt build, the pycnometer and ImageJ results were 98.01% and 98.20%, respectively. This closer alignment may be due to the concentration of small amorphous pores, which resulted in lower density in the stochastic spot melt build. Hardness values for the raster build and stochastic spot melt build, shown in Table 3.1.

Hardness values for the raster build and stochastic spot melt build, shown in Table 3.1. Observing Table 3.1 and the accompanying Figure 3.4, reveal that the raster build exhibits superior hardness when compared to the stochastic builds. This difference could be attributed to the internal porosity of the stochastic build, as pores can work as stress concentrators reducing hardness value.

This difference could be attributed to the internal porosity of the stochastic build, as pores can work as stress concentrators reducing hardness value. Table 3.1 Vickers Microhardness Value (HV1) for Raster and Stochastic Builds Build Test 1 Test 2 Test 3 Test 4 Test 5 Mean OverallStochastic (YZ) 403 426 419 427 422 419Raster(XZ) 433 436 417 439 433 432 431Stochastic (XZ) 425 426 416 394 419 416 418Raster(YZ) 428 428 430 428 434 430400420440Raster(XZ)Raster(YZ)Stochastic (XZ)Stochastic (YZ)Hardness Value (HV1)Scanning StrategiesVickers Hardness Value HV1 ScaleFigure 3.4 Vickers hardness value (HV1) for Raster and Stochastic builds Overall, during builds, the raster scanning strategy, associated with a slower and more controlled melting process, yielded components with a majority of columnar grains. This is likely due to the linear beam path allowing for a more ordered melting process and more effective gas expulsion.

This is likely due to the linear beam path allowing for a more ordered melting process and more effective gas expulsion. However, for stochastic melt scanning strategy, a higher Eas was required during the melting process. Even though this resulted in faster melting and produced fewer columnar grains in edges when paired with higher power and lower dwell time, it resulted in regions with higher porosity in the center. This porosity might be attributed to the stochastic nature of the melting strategy. During shorter dwell time or faster melting as described in Case I (Figure 3.2a), there could a rapid localized heating that quickly forms meltpool. Since the system does not have a controlled vacuum, even in a vacuum environment this might cause certain elements in metal powders to rapidly vaporize creating a localized vapor pressure within the meltpool.

This trapped vapor can affect the meltpool, leading to turbulence and forming465irregularly shaped pores as the vaporized material attempts to escape (Figure 3.2a). For Case II (Figure 3.2b), longer dwell time allowed the energy spread out over longer periods. This might have allowed the meltpools to maintain a uniform temperature reducing the rapid vaporization of the metal elements. With less vapor being produced, there could be a lower turbulence in meltpool, leading to a stable melting process even though the overall energy input remained same. Further investigation needs to be evaluated to make this above statement concrete.4 ConclusionIn summary, this study presented a comprehensive examination of the effects of raster and stochastic spot melt scanning strategies on the microstructure, density, and hardness of IN718 components produced via Electron Beam Powder Bed Fusion (EB-PBF).

The distinction in terms of melt pool dynamics, thermal gradient and solidification rates arising from these strategies resulted characteristics differences in microstructural features, defect formation, and overall part performance.A critical observation from this work is the need for careful balance of process parameters, particularly dwell time and power, to manage melting rates which affects solidification rate, prevent defects such as gas entrapped pores and keyholing, and ensure the production of high-quality parts in EB-PBF processes. Increased dwell time can alleviate the rapid melting related issues seen in stochastic spot melt, but at the risk of introducing additional energy-related defects.These findings underscore the importance of understanding the interplay between process parameters and material properties in additive manufacturing processes like EB-PBF.

They also serve to illustrate the potential of these techniques for tailoring microstructural properties to meet specific application requirements, highlighting the versatility and potential of EB-PBF in the broader field of materials engineering.Future research should continue to explore the effect of other process parameters and further optimize scanning strategies to achieve even more precise control over the resulting microstructure and properties. This will contribute to the ongoing evolution of EB-PBF, enhancing its utility and reliability for manufacturing a diverse range of high-performance components.AcknowledgementThe research presented here was conducted at The University of Texas at El Paso within the W.M. Keck Center for 3D Innovation (Keck Center), a 13,000-sq. ft. State-of-the-art additive manufacturing facility. The fund for this research was provided by the National Science Foundation -Grant Number 2117801.

ft. State-of-the-art additive manufacturing facility. The fund for this research was provided by the National Science Foundation -Grant Number 2117801. The author appreciates the assistance and opportunities extended by these institutions, which significantly contributed to the completion of this study.References[1] Q. C. Liu, Q. Guo, C. Li, Y. Mei, X. Zhou, and Y. Huang, "Recent progress on evaluation of precipitates in Inconel 718 superalloy," Acta Metall. Sin., pp. 1259-1266, 2016. Available: http://dx.doi.org/10.11900/0412.1961.2016.00290[2] R. G. Ding, Z. W. Huang, H. Y. Li, I. Mitchell, G. Baxter, and P. Bowen, "Electron microscopy study of direct laser deposited IN718," Mater. Charact., vol. 106, pp. 324-337, 2015. Available: https://doi.org/10.1016/j.matchar.2015.06.012466[3] J. R. Davis, Ed., Nickel, cobalt, and their alloys, ASM International.[4] A. Strondl, R. Fischer, G. Frommeyer, and A.

ploring Model-Based Engineering Concepts for Additive ManufacturingRobert R. Lipman, Jeremy S. McFarlaneSystems Integration Division, National Institute of Standards and Technology, Gaithersburg, MD 20899AbstractRobust geometry and tolerance representations are needed in additive manufacturing for precise part specification and interoperability with downstream activities such as manufacturing, inspection, and long-term archiving. A disconnection exists between process-independent part geometry and tolerances, and process-dependent information requirements for additive manufacturing. Existing and emerging standards for part geometry (ASTM AMF, 3MF, ISO 10303 STEP) and tolerances (ASME Y14) contain information related to the additive manufacturing process.

Existing and emerging standards for part geometry (ASTM AMF, 3MF, ISO 10303 STEP) and tolerances (ASME Y14) contain information related to the additive manufacturing process. Details of the standards will be discussed, how their use and improvement can benefit the additive manufacturing process, and their integration into the model-based engineering paradigm.1 – IntroductionThe trend for traditional, subtractive, manufacturing processes is towards a Model-Based Engineering (MBE) paradigm [1, 2]. The fundamental concept behind the adoption of MBE is to transition from communicating and exchanging part and manufacturing information in 2D drawings and documentation to using robust digital 3D product models and datasets. Traditionally, communicating part geometry and associated tolerances has been through drawings and associated documentation, methods that are well suited for human-consumption and interpretation.

Traditionally, communicating part geometry and associated tolerances has been through drawings and associated documentation, methods that are well suited for human-consumption and interpretation. This is not compatible with increased part complexity and digital manufacturing processes. Interpretation of drawings is a potentially error-prone process, to be avoided, when developing a process plan to program machine tools for manufacturing or coordinate measuring machines for inspection. The use of 3D product models and associated data exchange formats are better suited than drawings for digital applications for manufacturing, inspection, maintenance, and operations [3]. This drives the fundamental characteristic of MBE, a single dataset containing the digital 3D product model and other associated product definitions that support downstream applications throughout the lifecycle of a product. Additive manufacturing (AM) processes are not compatible with 2D drawings.

Additive manufacturing (AM) processes are not compatible with 2D drawings. AM processes align with the use of digital 3D product models and MBE practices.The U.S. defense industry has created a standard for MBE known as MIL-STD-31000A [4, 5]. The standard defines a Technical Data Package (TDP) used to describe the technical data used during the lifecycle of a product including “models, drawings, associated lists, specifications, standards, performance requirements, quality assurance provisions, software385documentation, and packaging details.” The core of a TDP includes the 3D product model of a part represented by a Computer-Aided Design (CAD) model and supplementary data exchange formats.

Currently, there is nothing specific to additive manufacturing in a TDP; however, the concept of the TDP being used as a lifecycle description of product information would be very beneficial for additive manufacturing.A fundamental characteristic of the MBE approach is to link part and product information across the stages of its lifecycle. An additive manufacturing process can be defined by eight phases of information [6]: (1) part design, (2) raw tessellated geometry, (3) final tessellated geometry, (4) build file, (5) machine data, (6) printed part, (7) finished part, and (8) validated part. Raw tessellated geometry (Phase 2), which approximates the precise geometry in the CAD model (Phase 1), is transferred to AM software. In the AM software, the tessellated geometry can be modified in a variety of ways to ensure manufacturability. The resulting tessellated geometry (Phase 3) is then sliced and prepared for printing (Phase 4).

The resulting tessellated geometry (Phase 3) is then sliced and prepared for printing (Phase 4). The slicing process can introduce further deviations from the original CAD geometry. Finally, the slices are interpreted by the machine (Phase 5) and the printed part can be inspected (Phase 8). There is also feedback between the phases. For example, results from the inspection can be used to modify the part design for a part that does not meet desired design specifications.Figure 1 shows a simplified flowchart of the design, printing, and inspection phases. Inspection is based on the tolerances specified with the original geometry in the CAD model. However, a disconnection exists between the design and inspection phases, indicated by the dotted line. The geometry of the printed part is based on an approximate tessellated geometry and its slices and not on the original geometry and tolerances from the CAD software.

The geometry of the printed part is based on an approximate tessellated geometry and its slices and not on the original geometry and tolerances from the CAD software. Consequently, there are two parallel independent representations of geometry. The first representation is the process-independent CAD geometry and tolerances. The second is a temporary representation used for printing that has transformed the CAD geometry into tessellated geometry and slices. The dual geometric representations and the disconnection between the design and inspection process create a significant challenge for an MBE process. The manufacturing and inspection processes would benefit if they were driven by a single geometric and tolerancing representation.386Figure 1: Flowchart of design, print, and inspection process2 - Representations for MBE in Additive ManufacturingRobust representations for geometry and tolerances are essential to the use of additive manufacturing information in MBE.

Those representations generally serve two purposes (1) to exchange information between both CAD and additive manufacturing software with downstream planning, printing, and inspection processes and (2) to provide a repository for an additive manufacturing product model. Many of the data exchange formats used for these representations satisfy only some of the needs of MBE. In this section we describe the relevant formats and standards for both geometric and tolerance representations.2.1 – Standards for Geometry RepresentationsPrevious research from the 1990’s and early 2000’s [7-10] evaluated the data requirements for additive manufacturing, available data exchange formats for geometry and slices, and proposed new formats. Some of the evaluated geometry formats include STL (Stereo Lithography) [11], STEP (Standard for Exchange of Product Data) [12], STH (Surface Triangles Hinted), and CFL (Cubital Facet List).

Some of the evaluated geometry formats include STL (Stereo Lithography) [11], STEP (Standard for Exchange of Product Data) [12], STH (Surface Triangles Hinted), and CFL (Cubital Facet List). Two new data exchange formats were also proposed: RPI (Rensselaer Polytechnic Institute) [13] and LMI (Layer Manufacturing Interface) [14] formats. Some of these formats have survived, while others have not. Regardless, many of the data exchange requirements these formats were intended to meet still exist. More recently, two new formats have been proposed: AMF (Additive Manufacturing Format) [15] and 3MF (3D Manufacturing Format) [16]. STEP has been improved relative to data exchange requirements for additive manufacturing. These formats and their applicability to meeting MBE requirements387in the context of additive manufacturing are described below.

These formats and their applicability to meeting MBE requirements387in the context of additive manufacturing are described below. A 2015 UK-funded project will also be assessing some of the same formats mentioned here, including AMF, STEP, and STEP- NC [17].STLThe STL (Stereo Lithography) format is the commonly used, de-facto standard for exchanging tessellated geometry from CAD or additive manufacturing software with other AM- related processes such as geometry slicing, support generation, and printing [11]. An STL file contains unstructured planar triangular facets (tessellations) defined by three vertices and a facet normal. As a result, an STL tessellation only approximates the exact CAD model geometry using these discrete triangular elements. Parameters, such as chord height, deviation, and angle tolerance, in the CAD software control the deviations between the exact and the tessellated geometry.

Parameters, such as chord height, deviation, and angle tolerance, in the CAD software control the deviations between the exact and the tessellated geometry. The tessellated geometry is unstructured in that every facet is defined completely independently of every other facet. There is no sharing of vertices between facets or defining edges and surfaces.STL files usually need to be repaired to ensure that they can be sliced and printed [18]. This entails making sure there are no missing triangles, collapsed edges, unshared edges, inverted normals, overlapping or intersecting triangles, and that the tessellated geometry is watertight. The STL tessellation can also be modified in other ways to optimize the printing process such as filling small holes, hollowing out solids, and thickening thin-walled parts. The STL format is still widely used despite acknowledged shortcomings.

The STL format is still widely used despite acknowledged shortcomings. These shortcomings, however, mean that STL does not satisfy the requirements for MBE.AMFThe Additive Manufacturing File (AMF) format is a data exchange standard developed by ASTM F42 committee on Additive Manufacturing Technologies [19, 20]. The AMF format is co-branded with the ISO TC 261 standards committee on additive manufacturing [15]. AMF is an XML-based file format defined by an XSD schema [21].The AMF format is intended as a replacement for STL. As such, it is a new means to transfer information between design software and additive manufacturing hardware. It has many features that are an improvement over the STL format.

As such, it is a new means to transfer information between design software and additive manufacturing hardware. It has many features that are an improvement over the STL format. These features include support for curved patches, recursive subdivision, multiple materials, graded materials, internal structures, material properties, colors, graphics, build placements, and metadata.Curved triangles are defined by normal vectors or edge tangents at each vertex of a planar triangle. Recursive subdivision further increases the tessellation accuracy by subdividing one triangle into four triangles.

Recursive subdivision further increases the tessellation accuracy by subdividing one triangle into four triangles. A highly accurate approximation of a sphere can be modeled starting with a coarse mesh of curved triangles and applying multiple levels of recursive subdivision.388This method generates a closer approximation of the precise geometry for a relatively coarse tessellation.Multiple and graded materials allow new materials to be created by specifying a mixing ratio of materials or a variation of their distribution within an object. The ratios or distributions of materials are defined by mathematical functions. Multiple materials can have different mechanical properties and when printed, according to their distribution, create new mechanical properties in specific regions of a part to satisfy design specifications.

Internal or external structures such as lattices or cooling fins can also be specified by a functional representation describing the distribution of material within a part.AMF also has the capability to specify colors or texture maps in a region or at vertices, graphic images such as logos, print orientation and layout in a build volume (constellation), and metadata strings. The standards documentation also lists future potential provisions for dimensional and geometric tolerances, surface roughness, support structures, functional representations, voxel representations, and surface texture.Although AMF is intended as a ‘printing’ format, the detailed model for additive manufacturing information satisfies only some of the needs of MBE. A single representation captures the intent of an additive manufacturing design. The AMF format also supports metadata that can be associated with, either the entire file or individual objects in the file.

The AMF format also supports metadata that can be associated with, either the entire file or individual objects in the file. Metadata includes the name and description of the object, revision number, an associated URL, producer information, object volume for verification purposes, a single tolerance value for the entire part, and material strength. The AMF format has been developed to be independent of model resolution and layer thickness and contains no information specific to any one additive manufacturing process.3MFThe 3MF Consortium introduced the 3D Manufacturing Format (3MF) in May 2015 as a new 3D printing format [16]. The consortium includes CAD and AM software vendors, 3D printer manufacturers, and 3D printing service bureaus. 3MF is designed to support interoperability between design software and 3D printing hardware. Documentation for 3MF includes two specifications and reference guides: (1) the 3MF core and (2) materials and properties.

Documentation for 3MF includes two specifications and reference guides: (1) the 3MF core and (2) materials and properties. 3MF is defined by an XSD schema and 3MF files are written in XML [21].The geometry representation for a 3MF object consists of a mesh defined by a list of vertices and triangles are defined by indices into the list of vertices. There can be multiple meshes in one object. Components and assemblies, such as a car body object and tire objects, can be defined each with its own transform to place the tires on the car body object. Objects can also reference a material name and color. The extension for materials and properties adds 2D texture coordinates, composite materials, and multi-properties to the 3MF core.389One of the primary purposes of 3MF is to send the required information to a 3D printer from design software directly. However, this purpose is not a concern for MBE.

However, this purpose is not a concern for MBE. In addition to the features described above, the 3MF file format includes other information packaged with the object geometry. That information includes a thumbnail image of the object, a print ticket, metadata, and a digital signature. Metadata includes information about the designer, copyright, license terms, creation and modification date, and build instructions.STEPSTEP (ISO 10303 –known informally as the STandard for Exchange of Product model data) [12, 22, 23] is a family of standards defining a methodology for describing product data throughout the life cycle of a product. Two STEP application protocols (AP) that have been widely implemented in CAD systems are AP203 known as ‘3D design of mechanical parts and assemblies’ [24] and AP214 known as ‘automotive mechanical design’ [25]. AP242, a new application protocol approved by ISO in 2014, is known as ‘managed model-based 3D engineering’ [26].

AP242, a new application protocol approved by ISO in 2014, is known as ‘managed model-based 3D engineering’ [26]. AP242 covers the scopes of both AP203 and AP214. Additionally, it contains many new or improved capabilities such as tolerances, kinematics, data quality, and tessellated geometry [27]. Each of these APs is specified by a schema in the EXPRESS language [28]. A ‘STEP file’, also known as a ‘Part 21 file’, refers to a file that is exported by CAD software [29].One purpose for STEP tessellated geometry [30] is to visualize large complex geometric models instead of rendering boundary representation geometry. In STEP, sets of triangulated faces are defined by a list of coordinates and indices referring to the coordinate list. Normal vectors are assigned to each triangle. Triangle strips or fans can be defined in addition to individual triangles. A coordinate list can contain all coordinates for an object or only those related to a set of triangulated faces.

Triangle strips or fans can be defined in addition to individual triangles. A coordinate list can contain all coordinates for an object or only those related to a set of triangulated faces. Multiple sets of triangulated faces can be grouped together to form a tessellated solid or shell. Multiple tessellated solids and shells can be grouped together to form a tessellated shape. Tessellated edges are defined by the two faces common to an edge and indices referring to the coordinate list for the endpoints of the edge. The definition of tessellated edges can be used to determine if the all of the triangles for an object define a watertight tessellation.STEP AP242 is perhaps the best suited for addressing some MBE issues in the context of additive manufacturing. In particular, a single STEP file can contain precise model geometry, tolerances, multiple tessellated geometry representations, and associations between them.

In particular, a single STEP file can contain precise model geometry, tolerances, multiple tessellated geometry representations, and associations between them. An important feature of STEP tessellated geometry is that a set of triangulated faces can refer to an exact geometric surface such as a plane, cylinder, or other boundary representations. Geometric and dimensional tolerances can also be associated with exact geometry and therefore with the tessellated geometry. Therefore, a set of triangulated faces can be directly inspected to ensure that they are within a tolerance value associated with them. This feature helps connect the disconnected processes shown in Figure 1.390Validation properties can be used to validate information in STEP files. For example, geometric validation properties are characteristics of solid and surface models, such as area, volume, and centroid that are written to a STEP file when it is exported from a CAD system.

For example, geometric validation properties are characteristics of solid and surface models, such as area, volume, and centroid that are written to a STEP file when it is exported from a CAD system. When the STEP file is imported to a receiving CAD system, that system can compute the same validation properties and compare them to the values from the originating system in the STEP file. If the computed validation properties are within an agreed tolerance to the original validation properties, then the exchange of geometric information has been validated. Validation properties have also been defined for tessellated geometry including the number of facets, surface area, and surface center point.Another essential feature of geometric representations is their use for long-term archiving and retrieval. This is an important aspect of model-based engineering in the aerospace, automotive, and defense industries.

This is an important aspect of model-based engineering in the aerospace, automotive, and defense industries. Information about the design of manufactured parts might need to be retrieved, decades from when it was first manufactured, to create replacement parts or understand the design parameters and rationale for a part. There is no guarantee that CAD models in proprietary formats will work with future versions of CAD software or computer systems. LOTAR is a consortium that develops, tests, publishes, and maintains standards for archiving and retrieval of digital data [31]. STEP application protocols are the basis for the standards being developed by LOTAR. Some of the LOTAR standards relate to explicit 3D CAD data and tolerances associated with CAD models. A LOTAR standard for additive manufacturing product data is also being considered.

Some of the LOTAR standards relate to explicit 3D CAD data and tolerances associated with CAD models. A LOTAR standard for additive manufacturing product data is also being considered. The LOTAR consortium also tests STEP translators in CAD software for conformance to the standards.The proposed scope of Edition 2 of AP242 includes higher-order curved triangles along with additive manufacturing specific information for the placement and orientation of parts in a build volume [32]. Past research also showed how additive manufacturing features such as heterogeneous materials [33, 34] and slices [35, 36] could be represented in STEP.

Past research also showed how additive manufacturing features such as heterogeneous materials [33, 34] and slices [35, 36] could be represented in STEP. As far back as 1995 STEP was being considered as a method to integrate design and solid freeform fabrication [37].STEP-NCSTEP-NC (Numerical Control) is a machine tool control language that (1) extends the ISO 10303 STEP standards with the ISO 14649 machining model and (2) includes geometric dimensioning and tolerancing information similar to AP242 [38-40]. The application protocol for STEP-NC is AP238. An AP238 file includes part geometry, features, tool characteristics, work plans, and machining parameters; however, it does not include tessellated geometry. The intention of STEP-NC is to control a machine tool directly from an AP238 file rather than generate G-code [41] to program a machine tool.Additive manufacturing is also being considered as an addition to STEP-NC.

The control of the movement and operation of a laser for powder-bed fusion or the motion of an extruder391head is considered the same as the control of a machine tool used for traditional subtractive machining operations. This development would also be of interest to control hybrid subtractive- additive manufacturing machines. Some of the characteristics of additive manufacturing being considered for a future version of AP238 include supports, lattices, skins, cores, layer strategy, and machine processes. A proof-of-concept has also been developed for representing extruder motion paths from a Common Layer Interface (CLI) file in a STEP AP238 file [42].STEP-NC supports use of machine tool control for MBE by having a single AP238 file that contains both geometry and a manufacturing process plan.

In the future, it may also include additive manufacturing information and tolerances.2.2 – Standards for Tolerance RepresentationsEqually important for MBE is the definition and communication of tolerances related to design geometry. Geometric dimensioning and tolerancing (GD&T) is a symbolic language for communicating engineering tolerances on manufactured parts. GD&T specifies the allowable deviation of features on a part so that the appropriate manufacturing and inspection processes can be used. The specification of GD&T on parts helps guarantee their operation in an assembly and controls how a part is measured to ensure that it meets the specified tolerances.The ASME Y14.5 and Y14.41 standards [43, 44] and ISO 1101 and 16792 standards [45, 46] define how GD&T is specified on 2D drawings and 3D models. None of the tolerancing standards define anything specific to additive manufacturing.

None of the tolerancing standards define anything specific to additive manufacturing. A new ASME committee Y14.46 was formed in 2015 to consider product definition practices for additive manufacturing, including tolerances. Research related to tolerances and additive manufacturing has considered part accuracy [47-49] and manufacturability of flatness and circularity tolerances [50-52].Figure 2 shows a typical CAD model with geometric and dimensional tolerances applied to surfaces, slots, and holes [53]. The presentation of the annotations on the drawing, as line segments, is meant to be human readable. The representation of the tolerances in CAD systems or other standard file formats is intended for consumption by downstream applications such as manufacturing and inspection.

The representation of the tolerances in CAD systems or other standard file formats is intended for consumption by downstream applications such as manufacturing and inspection. The annotation representation does not contain any information regarding their visual appearance.392Figure 2: CAD model with typical GD&T annotations [53].STEP AP242 contains information for the representation and presentation of ASME standards for tolerances. An AMF file can contain a single tolerance value related to the accuracy of the recursive subdivision. The AMF standard also lists dimensional and geometric tolerances as a possible future feature.Recent research considers tolerance specification issues for additive manufacturing [54]. Those issues are either related to the additive manufacturing process or highlighted by the capabilities of additive manufacturing.

Those issues are either related to the additive manufacturing process or highlighted by the capabilities of additive manufacturing. For example, the build orientation used in an additive manufacturing process can affect the part strength, requirements for support structures, surface roughness, build time, and other characteristics [55-59]. A designer might want to specify the build direction and acceptable deviations from the build direction to ensure that requirements for strength, tolerances, and surface roughness are met. The ASME Y14.5 standard defines a way to identify coordinate systems on 2D drawings. The coordinate system could be extended to specify the build orientation for each part in a single build. Specifying the layer thickness and orientation of laser scanning or deposition patterns has some similarity to the process-specific ASME Y14.37 standard for 2D drawings of composite parts [60].

Specifying the layer thickness and orientation of laser scanning or deposition patterns has some similarity to the process-specific ASME Y14.37 standard for 2D drawings of composite parts [60]. The composite standard contains a ply table of fiber orientations and layer thicknesses. The ply table concept could be393extended to specify layer thickness for printed parts and scan or deposition pattern directions. The definition of a grain direction in the process-specific ASME Y14.8 standard for casting, forgings, and molded parts [61] could also be adapted to represent scan or deposition pattern directions.3 – DiscussionTable 1 summarizes the geometric representations presented in section 2.1. Empty cells indicate that a geometric representation does not support a particular feature. The first row shows a ‘purpose’ for each of the representations. Clearly, STL, AMF, and 3MF are printing formats intended to transfer information from CAD and AM software to printing hardware.

The first row shows a ‘purpose’ for each of the representations. Clearly, STL, AMF, and 3MF are printing formats intended to transfer information from CAD and AM software to printing hardware. Both AMF and 3MF are intended to be replacements for STL with improved tessellated geometry and other features. Being supported by a consortium of software and hardware vendors, the 3MF format aims for a seamless printing process.The AMF format contains a lot more information to capture some of the unique features of additive manufacturing such as lattice structures and functionally graded materials. This positions the format as a ‘printing’ format with robust, additive-specific information. Some aspects of AMF put the burden of generating the final printed geometry on the slicing or printing software. Lattice structures are defined implicitly by a functional representation that has to be evaluated to generate the explicit geometry to be printed.

Lattice structures are defined implicitly by a functional representation that has to be evaluated to generate the explicit geometry to be printed. In addition to the AMF file format, ASTM has developed a standard terminology for coordinate systems used in additive manufacturing [62]. The standard can be used to specify the position and orientation of a part relative to the build platform.On the other hand, the STEP formats are product models and not printing formats. To ‘print’ a STEP file, the tessellated geometry can be converted to any of the ‘printing’ formats. The STEP file also maintains an association between precise geometry, tessellated geometry, and tolerances. In addition, STEP is the format used for long-term archiving and retrieval of product information.

In addition, STEP is the format used for long-term archiving and retrieval of product information. STEP validation properties can be used to verify geometry, precise and tessellated, and geometric and dimensional tolerances.There might need to be some middle ground between the ‘printing’ file formats and the robust product models. However, instead of a middle ground, additive manufacturing information might only be represented in a product model. The temporary ‘printing’ file formats are only used to send information to a printer. Consideration needs to be given to the additive manufacturing data exchange process, in the context of MBE, to maintain the integrity of the design geometry as modeled in CAD software and what is eventually printed, finished, and inspected regardless of which format is used.

Figure 3 shows a part with complex,395free-form surfaces and a part with multiple lattice structures and cooling fins. It is unlikely that traditionally specified tolerances or inspection methods can be applied to these and other features that are unique to additive manufacturing. For either part, even if a typical surface profile were applied to the entire part, it would be difficult to measure or inspect inaccessible surfaces. For the part on the left, a functional tolerance that measures the horizontal cross sectional area might be more appropriate. The area could be related to requirements for different loading conditions. For the part on the right, a tolerance might be applied to the outer boundary of each lattice section. New procedures will have to be developed to apply tolerances and inspect parts with inaccessible internal features such as cooling channels, embedded components, and transitions between heterogeneous materials.

J. Pratt, "Introduction to ISO 10303—the STEP standard for product data exchange," Journal of Computing and Information Science in Engineering, vol. 1, pp. 102-103, 2001. [23] A. B. Feeney, S. P. Frechette, and V. Srinivasan, "A Portrait of an ISO STEP Tolerancing Standard as an Enabler of Smart Manufacturing Systems," Journal of Computing and Information Science in Engineering, vol. 15, 2015. ISO 10303-203:2011, "Industrial automation systems and integration - Product data representation and exchange Part 203: Application protocol: Configuration controlled 3D design of mechanical parts and assemblies," for Standardization, Geneva, Switzerland. ISO 10303-214:2010, "Industrial automation systems and integration - Product data representation and exchange - Part 214: Application protocol: Core data for automotive mechanical design processes," International Organization for Standardization, Geneva, Switzerland.

ISO 10303-242:2014, "Industrial automation systems and integration - Product data representation and exchange - Part 242: Application protocol: Managed Model-based 3D Engineering," International Organization for Standardization, Geneva, Switzerland. [27] Development of a Convergent Modular STEP Application Protocol Based on AP 203 and AP 214: STEP AP 242 – Managed Model Based 3D Engineering, ASD Strategic Standardization Group, 2009, http://www.ap242.org/. ISO 10303-11:2004, "Industrial automation systems and integration - Product data representation and exchange - Part 11: Description methods: The EXPRESS language reference manual," International Organization for Standardization, Geneva, Switzerland. ISO 10303-21:2002, "Industrial automation systems and integration - Product data representation and exchange - Part 21: Implementation methods: Clear text encoding of the exchange structure," International Organization for Standardization, Geneva, Switzerland.

ISO 10303-42:2014, "Industrial automation systems and integration - Product data representation and exchange - Part 42: Integrated generic resource: Geometric and topological representation," International Organization for Standardization, Geneva, Switzerland.[24]International Organization[25][26][28][29][30][31] LOTAR - Long Term Archiving and Retrieval, http://www.lotar-international.org/. [32] Development of STEP AP242 Edition 2 - Managed Model Based 3D Engineering, ASDStrategic Standardization Group, 2014, http://www.asd-ssg.org/step-ap242-ed2.[33] L. Patil, D. Dutta, A. D. Bhatt, K. Jurrens, K. Lyons, M. Pratt, and R. D. Sriram, "A proposed standards-based approach for representing heterogeneous objects for layered manufacturing," Rapid Prototyping Journal, vol. 8, pp. 134-146, 2002.[34] M. Y. Zhou, "Modelling and representation of heterogeneous objects based on STEP for layered manufacturing," International Journal of Production Research, vol. 44, pp.

8, pp. 134-146, 2002.[34] M. Y. Zhou, "Modelling and representation of heterogeneous objects based on STEP for layered manufacturing," International Journal of Production Research, vol. 44, pp. 1297- 1311, 2006.398[35] H. Jee and B. Y. Lee, "Slicing STEP-based CAD models for CAD/RP interface," inProceedings of Solid Freeform Fabrication Symposium, 1999, pp. 171-178.[36] B. Starly, A. Lau, W. Sun, W. Lau, and T. Bradbury, "Direct slicing of STEP based NURBS models for layered manufacturing," Computer-Aided Design, vol. 37, pp. 387- 397, 2005.[37] C. R. Gilman and S. J. Rock, "The use of STEP to integrate design and solid freeformfabrication," in Proceedings of Solid Freeform Fabrication Symposium, 1995.[38] M. S. Ryou, H. S. Jee, W. H. Kwon, and Y. B. Bang, "Development of a data interface for rapid prototyping in STEP-NC," International Journal of Computer Integrated Manufacturing, vol. 19, pp. 614-626, 2006.[39] R. Bonnard, P. Mognol, and J.-Y.

[44] ASME Y14.41-2012, "Digital Product Definition Data Practices - Engineering Drawing and Related Documentation Practices," American Society of Mechanical Engineers, New York. ISO 1101:2012, "Geometrical product specifications (GPS) – Geometrical tolerancing – Tolerances of form, orientation, location, and run-out," International Organization for Standardization, Geneva, Switzerland. ISO 16792:2006, "Technical product documentation - Digital product definition data practices," International Organization for Standardization, Geneva, Switzerland. [47] D. Dimitrov, W. van Wijck, K. Schreve, and N. de Beer, "Investigating the achievable accuracy of three dimensional printing," Rapid Prototyping Journal, vol. 12, pp. 42-52, 2006.[45][46][48] G. D. Kim and Y. T.

Schreve, and N. de Beer, "Investigating the achievable accuracy of three dimensional printing," Rapid Prototyping Journal, vol. 12, pp. 42-52, 2006.[45][46][48] G. D. Kim and Y. T. Oh, "A benchmark study on rapid prototyping processes and machines: Quantitative comparisons of mechanical properties, accuracy, roughness, speed, and material cost," Proceedings of the Institution of Mechanical Engineers, Part B: Journal of Engineering Manufacture, vol. 222, pp. 201-215, 2008.[49] M. Bauza, S. P. Moylan, R. Panas, S. Burke, H. Martz, J. Taylor, P. Alexander, R. Knebel, R. Bhogaraju, and M. O'Connell, "Study of accuracy of parts produced using additive manufacturing," in 2014 ASPE Spring Topical Meeting: Dimensional Accuracy and Surface Finish in Additive Manufacturing, Berkeley, CA, 2014.[50] R. Arni and S. K. Gupta, "Manufacturability Analysis of Flatness Tolerances in Solid Freeform Fabrication," Journal of Mechanical Design, vol. 123, pp. 148-156, 1999. [51] T. Ollison and K.

Arni and S. K. Gupta, "Manufacturability Analysis of Flatness Tolerances in Solid Freeform Fabrication," Journal of Mechanical Design, vol. 123, pp. 148-156, 1999. [51] T. Ollison and K. Berisso, "Three-dimensional printing build variables that impactcylindricity," Journal of Industrial Technology, vol. 26, pp. 2-10, 2010.399Inc.,[52] K. Senthilkumaran, P. Pandey, and P. Rao, "Statistical modeling and minimization ofform error in SLS prototyping," Rapid Prototyping Journal, vol. 18, pp. 38-48, 2012.[53] R. Lipman and J. Lubell, "Conformance checking of PMI representation in CAD modelSTEP data exchange files," Computer-Aided Design, vol. 66, pp. 14-23, 2015.[54] G. Ameta, P. Witherell, S. Moylan, and R. Lipman, "Tolerance Specification and Related Issues for Additively Manufactured Products," in ASME 2015 International Design Engineering Technical Conferences & Computers and Information in Engineering Conference, IDETC/CIE 2015, Boston, Massachusetts, 2015.[55] K.

M. Phatak and S. Pande, "Optimum part orientation in Rapid Prototyping using genetic algorithm," Journal of Manufacturing Systems, vol. 31, pp. 395-402, 2012. [60] ASME Y14.37-2012, "Composite Part Drawings - Engineering Drawing and Related Documentation Practices," American Society of Mechanical Engineers, New York. [61] ASME Y14.8-2009, "Castings, Forgings, and Molded Parts - Engineering Drawing and Related Documentation Practices," American Society of Mechanical Engineers, New York. ISO/ASTM 52921:2013, "Standard Terminology for Additive Manufacturing - Coordinate Systems and Test Methodologies," American Society of Mechanical Engineers, New York.[62][63] P.Zelinski,AdditiveManufacturing’sManifoldBenefits,2012, http://www.mmsonline.com/blog/post/additive-manufacturings-manifold-benefits.[64] Additive Manufacturing and a New Revolutionin Design Engineering,2014, http://lockheedmartin.com/us/news/features/2014/additive-manufacturing.html.400.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperExploring the Manufacturability and Resistivity of Conductive Filament Used in Material Extrusion Additive Manufacturing Harry Gao and Nicholas A. Meisel Made By Design Laboratory School of Engineering Design, Technology, and Professional Programs, Penn StateABSTRACTAdditive manufacturing (AM) has the unique ability to build multifunctional parts with embedded electronics without the need for post-print assembly. However, many existing forms of multifunctional AM are not easily accessible to hobby-level users. Most hobby-level desktop 3D printers are only used with non-conductive filaments. Recently however, conductive filaments have become increasingly available for material extrusion desktop printers.

Most hobby-level desktop 3D printers are only used with non-conductive filaments. Recently however, conductive filaments have become increasingly available for material extrusion desktop printers. Ideally, the use of these filaments would allow circuitry to be printed simultaneously with the rest of the structure, enabling complex, inexpensive, multifunctional structures. However, the resistivity of conductive filament is significantly impacted by the geometry of the print and the printing parameters used in the build process. In this study, two types of commercially-available conductive filament were tested under a variety of parameters. It was found that print temperature, layer height, and orientation all significantly affect the resistivity in various ways. The knowledge from this research will allow users to design better multifunctional parts that have reduced resistivity.1.

The knowledge from this research will allow users to design better multifunctional parts that have reduced resistivity.1. INTRODUCTIONAdditive Manufacturing (AM, or 3D printing), is a process that forms complex three- dimensional parts by building layer by layer. Compared to traditional subtractive manufacturing processes, AM offers many unique advantages. With AM, complex geometries can be produced quickly with low tooling costs, reduced lead times, and rapid iteration [1]. The material extrusion process is one of the most popular and inexpensive form of AM. In this process, thermoplastic material is heated until it is past its melting point and then deposited onto a build platform. A fine nozzle is used to direct the deposition of material and a robotic structure controls the position of the nozzle to create the desired geometry [2].

A fine nozzle is used to direct the deposition of material and a robotic structure controls the position of the nozzle to create the desired geometry [2]. Material extrusion is simple, easy to learn, and highly accessible, making it popular among both hobbyists and industrial manufacturers.The layer-wise nature of AM allows abundant flexibility in the design process, including access to the entire internal structure of a design as it is being manufactured. This makes it possible to embed foreign components within the internal structure and then continue to print over them. By building components directly into the structure, the final product can be manufactured without the need to assemble, fasten, or glue anything afterwards. The ability to embed components leads to multifunctionality, where fully functional products are created by embedding different foreign components (particularly electronics) within the print.

The ability to embed components leads to multifunctionality, where fully functional products are created by embedding different foreign components (particularly electronics) within the print. Ideally, this would greatly shorten or even eliminate the time needed to assemble the product afterwards, speeding up production and reducing potential points of failure. However, many existing forms of printed electronics require expensive equipment and special inks, putting them out of reach for a large group of users.Recently however, carbon-based conductive filaments have been developed for use in desktop- scale material extrusion printers. The resistivity of these filaments have been measured by researchers under ideal circumstances [3], [4]. However, the material extrusion process can be1612used with a wide range of parameters and geometries, which may impact resistivity beyond what has been measured.

However, the material extrusion process can be1612used with a wide range of parameters and geometries, which may impact resistivity beyond what has been measured. For example, resistivity is usually measured based on a solid, homogeneous block of material, which is not what actually would be used in a multifunctional part. Existing research has not looked at the effect of factors such as nozzle temperature, layer height, geometry, or orientation on the resistivity of these conductive filaments. In addition, the effect of encasing conductive filament with non-conductive filament (as in an embedded component) has not been studied.

In addition, the effect of encasing conductive filament with non-conductive filament (as in an embedded component) has not been studied. This is important since multifunctional parts commonly make use of embedded filaments, and it is necessary to ensure that embedding does not increase resistivity of conductive traces.In this research study, two different types of conductive filament, Protopasta conductive PLA and BlackMagic conductive graphene, are tested under different printing parameters and geometries to evaluate their resistivity and manufacturability. In addition, the effect of embedding (via a dual-extrusion process) on resistivity is evaluated. Taking what is learned from the experimentation, the end goal is to determine guidelines that can allow potential users to better design for additive manufacturing when using conductive filaments.2. EXISTING RESEARCH OF MULTIFUNCTIONAL ELECTRONICS IN AM Related research has been divided into three sections.

EXISTING RESEARCH OF MULTIFUNCTIONAL ELECTRONICS IN AM Related research has been divided into three sections. Section 2.1 discusses the possibilities of multifunctional printing. Section 2.2 discusses direct write technology with conductive materials. Section 2.3 discusses recent studies into material extrusion conductive filaments.2.1 Multifunctional AMSince its introduction, AM has offered many advantages in the way components can be designed and made. The freedom of design for additive manufacturing (DfAM) and the availability of an ever-expanding library of materials has led to the growth of multifunctional AM. While standard AM uses just one material, the idea behind multifunctional printing is combining different types of materials - particularly electronics - so that functional components can be produced without the need to assemble any parts after the build process.

Some recent work has explored the use of AM with conductive materials to see what can be done with current technology.Aguilera et al. [5] designed and built a printed electric motor, taking advantage of the embedding capability of AM. Components such as the magnets, controllers, and wiring were embedded directly into the structure of the motor. The end result was a custom-built, fully functional motor that did not require any post-printing assembly. Similarly, MacDonald et al. [6] compared the use of AM (via direct write conductive inks and stereolithography) to traditional injection molding for the purposes of prototyping and manufacturing an electronic gaming die. The researchers found that the time to market could be significantly improved using AM compared to traditional manufacturing. They estimated that the additive method was at least 400% faster than the traditional method.

They estimated that the additive method was at least 400% faster than the traditional method. By harnessing the capability to direct write and embed electronic circuitry into the products, they could be evaluated for fit, form, and function simultaneously. Klomp [4] explored the use and applications of the material extrusion method with conductive filaments. The researchers printed a capacitive touch wheel and a capacitive touch sphere using dual extrusion and conductive filament. By connecting the capacitive sensors to Arduinos, they were able to use them to control LED lights, music, and even video games.

By connecting the capacitive sensors to Arduinos, they were able to use them to control LED lights, music, and even video games. While these are only a few examples1613of how multifunctional AM has been utilized in product design, they nevertheless capture the possibility offered by multifunctional embedding or direct write.2.2 Direct WriteDirect writing describes processes that can deposit thin layers of conductive materials at very high accuracy to support the creation of multifunctional structures. Hon [7] discusses many different methods of direct writing, including inkjet, laser, and mechanical systems. Direct writing is usually but not always related to electronic circuitry and conductive materials. As an additive process, direct writing offers many similar advantages such as reducing waste material, enabling greater design flexibility, and speeding up the manufacturing process [8].

As an additive process, direct writing offers many similar advantages such as reducing waste material, enabling greater design flexibility, and speeding up the manufacturing process [8]. There is significant overlap between technologies used for direct writing and for additive manufacturing. For example, extrusion-type direct write machines are very similar in basic function to a common material extrusion 3D printer. Inkjet and aerosol jet direct writing is similar to the polymer jetting method of additive manufacturing. However, compared to AM (especially material extrusion), most forms of direct writing are designed for specialized purposes and not easily accessible to an average user.There has also been significant research into the impact of design and processing on the conductivity of direct-write traces. For example, Perez [9] deposited silver conductive ink and embedded it into various geometries during the material jetting process, and the resistivity was observed.

For example, Perez [9] deposited silver conductive ink and embedded it into various geometries during the material jetting process, and the resistivity was observed. Encapsulating the conductive ink with non-conductive material was found to significantly increase the resistivity of traces compared to fully-exposed conductive ink. The researchers concluded that, due to the high resistivity of their sample traces, they would not be well suited for high power applications. However, they demonstrated that the method could be used for strain gauge and capacitive sensor applications.2.3 Conductive filamentCurrently, multifunctional printing and direct writing are out of reach of many hobby-level users, due to the need for expensive conductive inks and specialized equipment. In recent years, researchers and manufacturers have been working to develop carbon-composite filaments that would be capable of being used with common desktop-scale material extrusion printers.

In recent years, researchers and manufacturers have been working to develop carbon-composite filaments that would be capable of being used with common desktop-scale material extrusion printers. Since material extrusion is by far the most widely accessible form of AM, this would allow multifunctional printing to reach a much wider audience. Of course, the filament must not only be accessible but also sufficiently conductive. While matching the resistivity of direct write inks would be ideal, higher resistivity can still be useful for sensors and gauges, as stated by Perez [9].In general, conductive filaments are created by adding carbon to standard filament in order to improve its properties. By mixing small quantities of these materials into standard filament, the filament will takes on some of the properties of the carbon material, but can still be heated and extruded normally. Ibrahim et al.

Ibrahim et al. [3] performed an initial study on the resistivity of Protopasta conductive PLA under different lengths and volumes, and compared its resistivity to a conductive ABS filament. The researchers tested the resistivity of different lengths of the filament (before extrusion), as well as the resistivity of single strands of extruded filament. In both cases, the resistivity was found to be abnormally high within the first 10 mm of filament, quickly dropping down to a more consistent level around 20 mm. They found the resistivity of the extruded filament to be around 0.25 ohm-meters, which correlates with ProtoPlant's claims. Klomp [4] tested several1614different conductive filaments, including Protopasta conductive PLA. They observed a resistivity of 0.46 ohm-m along the layers of a 3D-printed cube, and 0.84 ohm-m across the layers. Protopasta was found to be the best of the filaments they tested, being both the least resistive and the easiest to print.

Protopasta was found to be the best of the filaments they tested, being both the least resistive and the easiest to print. The numbers they observed are somewhat higher than Ibrahim and Protopasta’s claims.As this review of literature has shown, existing work has emphasized the mechanical properties and some potential applications of conductive filaments. However, the aspect of design for additive manufacturing with these new filaments has largely been ignored. This aspect is particularly important since it is unknown how the resistivity of these filaments can vary based on different conditions, including print parameters and part geometries. The goal of this research is to try to address this gap in knowledge. With a better understanding of the effects of different parameters, users will be able to design and print multifunctional parts with resistivity that better matches the designer’s intended application.3.

EXPERIMENTAL APPROACHThere are three primary objectives of this research study. The first is to investigate the effects of print parameters on the resistivity of conductive filaments, including nozzle temperature, layer height, extrusion speed, and travel speed. The second objective is to investigate the effects of print geometry on resistivity. This includes part orientation as well as single strands compared to printed blocks. The third objective is to investigate the effect of enclosing conductive material using a dual extrusion process. The work presented will determine which parameters and geometries will produce the lowest resistivity in the printed traces and will act as a guide for designers looking to maximize the effectiveness of these traces in low-cost printed electronics.3.1 Materials and equipmentThe printer used in this experiment is a Lulzbot TAZ5 fitted with a dual extrusion head.

It was chosen for this experiment due to its versatility and compatibility with a dual extrusion head as well as custom filament materials [10]. The printer was controlled with Cura slicing software, which contains pre-made “quickprint” parameters for hundreds of unique filaments, one of which is Protopasta conductive PLA [11].An important difference between material extrusion and other forms of manufacturing is that parts made with material extrusion exhibit significant anisotropic properties affected by the orientation of the layers compared to the mechanical or electrical load [12]. As such, build orientation nomenclature is essential. According to ASTM standards [13], part orientation is named with the longest dimension first and the second-longest dimension second. X and Y are the two horizontal dimensions and Z is used for the vertical dimension.

X and Y are the two horizontal dimensions and Z is used for the vertical dimension. As shown in Figure 1, the XY orientation refers to a part lying flat on the build plate, the ZX orientation refers to a part standing vertically like a column, and the XZ orientation refers to a part lying horizontally on its edge. The authors use this standardized naming approach throughout the paper.1615Figure 1. Specimens on the build platform in various orientations: (a) XY, (b) ZX, and (c) XZ.The two conductive filaments described previously were tested: Protopasta conductive PLA and BlackMagic conductive graphene filament. Protoplant claims that its conductive PLA filament can achieve up to 0.3 ohm-m in-plane with the layers and 1.15 ohm-m perpendicular to the layers (in the ZX orientation) [14]. As previously discussed, Ibrahim et al. [3] were also able to observe a resistivity of approximately 0.25 ohm-m. The second conductive filament is Conductive Graphene Filament made by BlackMagic3D.

As previously discussed, Ibrahim et al. [3] were also able to observe a resistivity of approximately 0.25 ohm-m. The second conductive filament is Conductive Graphene Filament made by BlackMagic3D. The company claims that the filament has a volume resistivity as low as 0.006 ohm-m [15].3.2 Selection of independent variablesFour independent variables were chosen to be analyzed in the initial set of experiments: Print temperature, layer height, extrusion speed, and travel speed. Once some consistent trends were established with the initial horizontal (XY) experiments, the vertical (ZX) experiments only looked at layer height as the independent variable. Finally, the independent variable for the dual extrusion experiments was how much of the conductive sample’s cross-sectional perimeter was enclosed.Five levels were chosen for each independent variable, summarized in Table 1 below. The levels were chosen based on recommended manufacturer settings and common 3D printer settings.

The levels were chosen based on recommended manufacturer settings and common 3D printer settings. In addition, manufacturability was used to determine the range of parameters. For example, test samples would not print reliably below 200ºC or above 240ºC. At temperatures below 200ºC, the filament was not hot enough to extrude, while at temperatures above 240ºC, the extruded filament would flow excessively and would not adhere to the intended structure.Table 1. Table of independent variables and parametersTemperature (C) 200 210 220 230 240Layer Height (mm) 0.2 0.3 0.4 0.5 0.6Extrusion Speed (mm/s) 10 30 50 70 90Travel Speed (mm/s) 140 160 180 200 2203.2 Single Strand SamplesThe first part of the experimentation consisted of testing single strands of filament, for both Conductive PLA and graphene materials.

The samples were printed with nominal dimensions of 0.5 mm x 0.5 mm x 30 mm; however, due to the flexibility of extruded material before cooling and the use of different layer heights, the actual dimensions of the samples varied. For example, the strand printed at 0.2 mm layer height had a height of 0.2 mm and a width of 0.8 mm.1616Meanwhile, a layer height of 0.5 mm would result in dimensions closer to 0.5 x 0.5 mm. Among samples with the same nominal dimensions, the physical dimensions varied by about ±0.05 mm. The cross-sectional width and height of the samples were measured using a Neiko digital caliper with an accuracy of 0.02 mm.Figure 2. A single strand sample being tested for resistivityThe resistance of the samples were tested using a Hewlett-Packard 972A handheld multimeter. The accuracy for ohm measurements is 0.2%. The hook leads were pushed through a block of foam to maintain a constant 20 mm distance.

The accuracy for ohm measurements is 0.2%. The hook leads were pushed through a block of foam to maintain a constant 20 mm distance. The distance between the leads was confirmed via digital caliper. The resistivity was then determined by using the equation𝑅𝑒𝑠𝑖𝑠𝑡𝑖𝑣𝑖𝑡𝑦 (𝑜ℎ𝑚 ∗ 𝑚) =𝑅𝑒𝑠𝑖𝑠𝑡𝑎𝑛𝑐𝑒 (𝑜ℎ𝑚) ∗ 𝑤𝑖𝑑𝑡ℎ (𝑚) ∗ ℎ𝑒𝑖𝑔ℎ𝑡 (𝑚) 𝑙𝑒𝑛𝑔𝑡ℎ (𝑚)The conductivity could then be found by taking the inverse of the resistivity. To analyze trends across multiple changing variables, a Response Surface Methodology (RSM) was used. Minitab was used to generate a full response surface design, consisting of 31 runs, for the horizontal single- strand testing. For each run, three samples were printed using the parameters defined in the run.3.3 Horizontal Block SamplesThe second geometry was a thin, rectangular block, designed to test the layer-on-layer and road-to-road effects of printed filament. The dimensions of the block were 1.5 mm x 1.5 mm x 30 mm.

The dimensions of the block were 1.5 mm x 1.5 mm x 30 mm. There was a difference of up to ±0.2 mm in the measured dimensions compared to the nominal dimensions. The samples were attached to the multimeter using alligator-type leads, shown in Figure 3. This method was used to measure the resistivity along the layers (in the XY orientation). The same response surface design used for the single strand testing was used again here. This included 31 different runs and 3 samples tested for each run.Figure 3. A horizontal block sample being tested for resistivity3.4 Vertical Block SamplesThe third geometry was similar to the second, but this time printed in the vertical (ZX) orientation so that the layers were stacked on top of each other. However, printing in the ZX1617orientation is much more unstable since the base size is very small and tall samples tend to wobble when nearing the top of the print. Because of this, 2 x 2 mm dimensions were chosen for the vertical sample.

Because of this, 2 x 2 mm dimensions were chosen for the vertical sample. Figure 4 shows an example of a 2 x 2 mm vertical sample. The samples were measured in the same fashion, but due to the different orientation, the resistivity across layers (ZX orientation) was now being measured.Figure 4. A 2 x 2 mm vertical sample with a layer height of 0.5 mmAfter a set of optimal build parameters had been determined from the horizontal testing, they could be used to narrow down the experimental variables for this geometry experimentation. Print temperature was held constant at 240℃, extrusion speed was held at 50 mm/s, and travel speed was held at 180 mm/s for the conductive PLA material. For the graphene filament, the print temperature was held constant at 230℃, extrusion speed was held at 30 mm/s, and travel speed was held at 180 mm/s.

For the graphene filament, the print temperature was held constant at 230℃, extrusion speed was held at 30 mm/s, and travel speed was held at 180 mm/s. The only independent variable for the ZX oriented specimens was the layer height, which again ranged from 0.2 mm to 0.6 mm in 0.1 mm increments.3.4 Dual extrusion samplesFinally, dual extrusion samples were tested in the horizontal (XY) and vertical (ZX) orientations. In the horizontal sample, the central non-conductive section was exactly 20 mm long and the ends of the conductive section were left exposed. This was done so that the alligator clips could be attached right against the central section, ensuring a constant distance for measurement.Figure 5. Horizontal and vertical dual extrusion samples.In the vertical orientation, the sample geometry had to be changed. In order to print the previous sample in the vertical orientation, the non-conductive sheath would have to be printed in the air.

In order to print the previous sample in the vertical orientation, the non-conductive sheath would have to be printed in the air. The geometry was too small to attach support material below the sheath. Printing would cause the sheath section to sag during the build process, making measurement inaccurate. Instead, a new sample and method of measurement was chosen. The sample and sheath were both printed to be161820 mm tall. To measure the resistance, copper tape was attached to the top and bottom ends of the sample, which were then connected to the multimeter via alligator clips. For the dual extrusion testing, the only independent variable was whether or not the conductive sample was encased by a non-conductive sheath. The samples were printed using the same parameters discussed previously. All horizontal samples were printed at 0.2 mm layer height, and all vertical samples were printed at 0.6 mm layer height.4.

The samples were printed using the same parameters discussed previously. All horizontal samples were printed at 0.2 mm layer height, and all vertical samples were printed at 0.6 mm layer height.4. RESULTS AND ANALYSISAs discussed in Section 3, response surface methodology (RSM) was used to analyze the results of the experiments. The RSM method fits a second-order polynomial to the model. In a RSM analysis of variance, the linear p-value determines the significance of a linear trend. The square p-value determines the significance of a parabolic trend. A parabolic trend suggests that the variable has a maxima or minima, unlike a simple linear relationship. The parabolic trend can be independent of the linear trend, so there does not need to be a significant linear p-value in order for the square p-value to be valid. In addition, the RSM analysis of variance also determines the relationship of interaction effects.

In addition, the RSM analysis of variance also determines the relationship of interaction effects. A statistically significant two-way interaction effect can be confirmed as long as one of the individual main effects is also significant.4.1 Single Strand PLA SamplesAssuming that a p-value of 0.05 or less denotes statistical significance, it can be seen in Table 2 that there is a significant linear main effect for both nozzle temperature and extrusion speed with respect to the resistivity of the sample. There is also a significant parabolic main effect for temperature. This indicates that the relationship between temperature and resistivity can be described as a parabola. By looking at the following main effects plot in Figure 6, we can see the trends suggested by the data.Table 2. Analysis of Variance for the Single-Strand Conductive PLA samples.Variable Temperature Extrusion Speed Temperature*Temperaturep-value <0.001 0.047 0.007Figure 6.

Analysis of Variance for the Single-Strand Conductive PLA samples.Variable Temperature Extrusion Speed Temperature*Temperaturep-value <0.001 0.047 0.007Figure 6. Main effects plot for the single-strand conductive PLA samples.The analysis of variance says that the temperature plot on the far left of Figure 6 is statistically significant, both in the linear and parabolic relationships. The resistivity is at a minimum around1619230ºC. This correlates with Protoplant's recommended print temperature. The layer height and travel speed plots are not statistically significant. For extrusion speed, the analysis of variance says that the linear trend is significant, but the parabolic trend is not. In other words, resistivity generally increases with greater extrusion speed, but there is probably not a significant local minimum as suggested by the plot.It makes intuitive sense for higher temperature to result in lower resistivity. Zhang et al.

Zhang et al. [16] suggested that the extrusion process affects resistivity by orienting the conductive particles. A higher printing temperature increases the filament viscosity [17], which may allow the particles to flow more freely and better distribute themselves within the filament instead of getting agglomerating. An extrusion speed that is too high may lead to inaccurate deposition [17], which could cause the conductive particles to be further apart than desired. Conversely, an extrusion speed that is too low may cause clogging and inconsistent deposition. In this case, the layer height is not significant because only one layer was printed; its largest effect was changing the dimensions of the single layer.

In this case, the layer height is not significant because only one layer was printed; its largest effect was changing the dimensions of the single layer. Also, travel speed is not significant since only a single strand was printed, so there are no actual travel movements.The measured resistivity of the single-strand filament ranges from 0.045 ohm-m in the best- case scenario (print temperature at 230°C) to 0.055 ohm-m at a print temperature of 200°C. In comparison, Ibrahim [3] observed a value of 0.25 ohm-m for single-strand extruded filament, with print temperature in the range of 210ºC to 230ºC. The observed resistivity was around 80% lower than Ibrahim’s data. This may be due to the use of a different method of measurement.4.2 Single Strand Graphene SamplesFor the graphene samples, only temperature was found to have a statistically significant effect. Again, temperature had a parabolic relationship with a minimum resistivity at 230ºC.

Again, temperature had a parabolic relationship with a minimum resistivity at 230ºC. The overall resistivity was only a small amount lower than the Conductive PLA, with a range of about 0.01 ohm-m to 0.05 ohm-m.Table 3. Analysis of Variance for the Single-Strand Graphene samples.Variable Temperature Temperature*Temperaturep-value <0.001 0.039Figure 7. Main Effects Plot for Single-Strand Graphene samples1620This measured resistivity is approximately 2-8 times larger than the resistivity claimed by BlackMagic for their filament. This could be due to a number of factors. First, BlackMagic did not specify whether their value was measured based on a printed sample or if it was based on a pure block of graphene. Second, when measuring very low resistances, the magnitude of any possible errors are magnified.

Second, when measuring very low resistances, the magnitude of any possible errors are magnified. While the actual resistances that were being measured should be well within the capability of the multimeter, any sort of error caused by an imperfect connection to the sample or an incorrect caliper measurement could have a significant impact on the result. For instance, the measurement error of the calipers is ±0.02 mm, which is up to 5% of a single dimensional measurement. When combined with the 3 dimensions, and the fact that an error in length may also affect the measured resistance, the effect of caliper error could be 15-20%.4.3 Horizontal (XY) Block PLA SamplesThe analysis of variance for the PLA horizontal block samples shows that both temperature and layer height were statistically significant based on a linear relationship. However, the parabolic relationships were not significant.

However, the parabolic relationships were not significant. This suggests that higher temperature (240ºC) and lower layer height (0.2 mm) tend to reduce resistivity in the horizontal (XY) orientation. The effect of extrusion speed was found to be not significant in the linear but significant in the parabolic. This suggests that there is an ideal extrusion speed at approximately 50 mm/s, and going too far below or above this speed will cause resistivity to increase. The effect of travel speed was again not found to be significant.Table 4. Analysis of Variance for the Horizontal Block Conductive PLA samples.Variable Temperature Layer Height Extrusion Speed*Extrusion Speed Temperature*Extrusion Speedp-value <0.001 0.002 0.009 0.029Figure 8. Main Effects Plot for the Horizontal Block Conductive PLA Samples.The analysis of variance also shows that there is a statistically significant interaction between temperature and extrusion speed.

Main Effects Plot for the Horizontal Block Conductive PLA Samples.The analysis of variance also shows that there is a statistically significant interaction between temperature and extrusion speed. A higher extrusion speed is more desirable as the print temperature increases. While the overall minimum resistivity seems to occur at an extrusion speed in the range of 40-50 mm/s, the minimum resistivity for a print temperature of 240°C seems to occur at an extrusion speed between 50 and 70 mm/s. The temperature and layer height results align with the original hypothesis that higher temperature and lower layer height would reduce resistivity. As stated earlier, a higher temperature may allow the conductive particles to flow more1621freely. The effect of lower layer height is also as expected. Lower layer height means that the dimensions of the block are filled more completely and there are fewer gaps between printed roads.

The effect of lower layer height is also as expected. Lower layer height means that the dimensions of the block are filled more completely and there are fewer gaps between printed roads. This should result in less contact resistance as the current flows in the direction of the layers.The observed resistivity for each of the block samples is much higher than the single-strand samples, which makes sense since contact resistance now plays a much greater role in overall resistivity. Based on the contour plot, it appears that an ideal sample printed at 240ºC with 0.2 mm layer height and 50 mm/s extrusion speed should have a resistivity of close to 0.1 ohm-m. In comparison, the manufacturer claims a resistivity of 0.25 ohm-m for printed parts in the XY orientation [14], and Klomp observed a resistivity of 0.46 ohm-m in the XY orientation [4]. The resistivity observed this study is about 60-80% lower. One potential explanation for this discrepancy is the difference in sample geometry.

The resistivity observed this study is about 60-80% lower. One potential explanation for this discrepancy is the difference in sample geometry. The standard sample geometry for volume resistivity testing is a 10 mm cube, and this is what was used by both the manufacturer and Klomp. However, Ibrahim [3] found that the resistivity of single-strand conductive PLA was abnormally high in the first 10 mm, dropping down to a consistent value after 20 mm. Further research should be done to see if this may have an effect on printed samples as well.4.4 Horizontal (XY) Block Graphene SamplesFor the graphene samples, only temperature and layer height were found to be significant, and this time temperature had a parabolic relationship. The temperature has a minimum at about 230ºC, which correlates well with BlackMagic's recommended print temperature of 220°C. This time the extrusion speed was found to have very little significance at all.Table 5.

This time the extrusion speed was found to have very little significance at all.Table 5. Analysis of Variance for the Horizontal Block Graphene samples.Variable Temperature Layer Height Temperature*Temperaturep-value 0.018 0.012 0.040Figure 9. Main Effects Plot for Graphene Horizontal Block samplesFrom Figure 9, the overall resistivity ranges from just under 0.06 ohm-m to 0.085 ohm-m, which is two orders of magnitude higher than the manufacturer’s claimed resistivity. While this seems significant, as noted before, the manufacturer’s claimed resistivity may not have been measured based on a printed part. BlackMagic readily admits that "the resistivity will change depending on your print" so a higher resistivity than the ideal is to be expected [15].

BlackMagic readily admits that "the resistivity will change depending on your print" so a higher resistivity than the ideal is to be expected [15]. In this case,1622it is likely that the contact resistances between the printed layers and strands are preventing the current from flowing as quickly as it could in a homogeneous block of material.4.5 Vertical (ZX) Block PLA SamplesAfter several rounds of testing, the effect of print temperature was found to relatively consistent, with a temperature of 230-240ºC being the best for most samples. The effect of extrusion speed was also relatively consistent, with an optimal extrusion speed at 50 mm/s for PLA. The ideal numbers also generally match up well with the parameters recommended by the manufacturers. Travel speed has been found to be insignificant in all cases so far. Due to these findings, for the vertical (ZX) block samples, experimentation focused on the most important variable, layer height.

Travel speed has been found to be insignificant in all cases so far. Due to these findings, for the vertical (ZX) block samples, experimentation focused on the most important variable, layer height. Each layer height was tested three times, and the results were averaged. A one-way ANOVA was performed to determine the statistical significance of the effect of layer height on ZX block specimens; the calculated p-value was less than 0.001.0.550.5)0.45m * m h o ( y t i v i t s i s e R0.40.350.30.250.20.150.10.20.30.40.50.6Layer Height ( mm)Figure 10. Plot of resistivity vs. layer height for the vertical conductive PLA samples.The resistivity results of the vertical samples were contrary to the original hypothesis. Instead of lower layer heights causing lower resistivity, Figure 10 shows that higher layer heights result in lower resistivity.

Instead of lower layer heights causing lower resistivity, Figure 10 shows that higher layer heights result in lower resistivity. Of course, this is for the vertical orientation only - in the horizontal, lower layer heights still produce less resistivity. This result is surprising at first, but makes sense on closer examination. In the horizontal orientation, the current was flowing along the layers - as if flowing through a wire. However, in the vertical orientation, the current is flowing across the layers. A low layer height means a much larger number of interfaces that the current must cross, and each one has contact resistances. A higher layer height means much fewer interfaces, and therefore less resistance. See Figure 11 for a visualization. In the space of one 0.6 mm layer, there could be three 0.2 mm layers.

A higher layer height means much fewer interfaces, and therefore less resistance. See Figure 11 for a visualization. In the space of one 0.6 mm layer, there could be three 0.2 mm layers. In the 0.2 mm sample, the current would have to cross three different layers, while in the 0.6 mm sample, the current can flow unimpeded across a homogeneous single layer.1623Figure 11. A schematic comparing interfaces for 0.2 mm layer height with 0.6 mm layer height.As the data shows, the 0.4 mm layer height has almost the same resistivity as the 0.6 mm layer height in the vertical orientation. However, it has much lower resistivity in the horizontal orientation. Therefore, if a component needs to be printed using conductive filament in both horizontal and vertical orientations, then 0.4 mm layer height would be a well-balanced compromise choice.

Therefore, if a component needs to be printed using conductive filament in both horizontal and vertical orientations, then 0.4 mm layer height would be a well-balanced compromise choice. Protoplant originally claimed a resistance of 1.15 ohm-m when measuring resistivity across the layers [14], and Klomp observed a resistance of 0.84 ohm-m [4]. These results show resistivity that is around 60-80% lower. The resistivity would depend on the layer height that Protoplant and Klomp used when testing, which is unknown. If they used 0.2 mm or lower, then the results differ by the same margin as the horizontal samples.4.6 Vertical (ZX) Block Graphene SamplesFor the graphene samples, a similar decreasing trend can be seen in Figure 12, with a p-value of 0.3745. However, since the overall resistivity of the graphene samples is much lower (while the standard deviation was similar), the results cannot be considered statistically significant.

However, since the overall resistivity of the graphene samples is much lower (while the standard deviation was similar), the results cannot be considered statistically significant. Based on the behavior of the conductive PLA, it seems likely that graphene should also exhibit less resistivity with higher layer height. However, more accurate testing must be done in order to confirm this.0.2)m * m h o ( y t i v i t s i s e R0.180.160.140.120.10.20.30.40.50.6Layer Height ( mm)Figure 12. Plot of resistivity vs. layer height for the vertical graphene samples.4.7 Dual Extrusion SamplesThe dual extrusion experiments were conducted to see if enclosing the conductive filament with regular non-conductive filament would affect its resistivity. In an actual multifunctional part, most conductive filament would be enclosed, so this is an important factor in overall resistivity.

CONCLUSIONS AND FUTURE WORKThe goal of this research study was to analyze the effect of print parameters such as temperature, layer height, extrusion speed, and travel speed on the resistivity of conductive filaments. The second major goal was to analyze the effects of on resistivity. The third major goal was to determine whether dual extrusion would affect resistivity. Two different types of commercially-available conductive filament were selected, tested, and compared. Various samples were printed under a range of different parameters and geometries, and their resistivities were analyzed using response surface methodology. Based on these tests, the authors conclude that:Generally, higher print temperature correlates to lower resistivity, up to certain limits. Conductive PLA does best at around 240°C, and graphene does best at around 230°C.

Conductive PLA does best at around 240°C, and graphene does best at around 230°C. Excessively high nozzle temperatures cause unreliable extrusion.Smaller layer height reduces resistivity when measuring along the filament (horizontal orientation). When measuring across the filament (vertical orientation), smaller layer height increases resistivity. A good balance can be struck at 0.4 mm layer height if conductivity in both orientations is required.The best extrusion speed for conductive PLA is around 50 mm/s. Extrusion speed does not appear to have a significant effect on the resistivity for graphene.Travel speed does not appear to have a significant effect on resistivity at all.  Enclosing a conductive trace within standard non-conductive filament does not appear to impact its resistivity.To expand the breadth of this research, future work should focus on studying applications of this technology.

The knowledge of print parameters and the effects of different geometries can be applied to the design of printed electronics, such as capacitive sensors and strain gages. Future work could study how to best use conductive material extrusion printing to produce low-cost, customizable electronic components. In addition, future work could look into the chemical composition of conductive filaments to gain a more concrete understanding of why the resistivity changes under different conditions, which may lead to more useful and versatile conductive filaments in the future. Another potential area of research would be studying a wider range of conductive filaments. In addition to conductive PLA and graphene, there are a range of other commercially available conductive filaments, including filaments that are loaded with silver or copper particles.16256. REFERENCES [1]P. Isanaka and F.

PLORING THE MECHANICAL PERFORMANCE AND MATERIAL STRUCTURES OF INTEGRATED ELECTRICAL CIRCUITS WITHIN SOLID STATE METAL ADDITIVE MANUFACTURING MATRICESJ. Li*, T. Monaghan, A. Bournias-Varotsis, S. Masurtschak, R. J. Friel, and R. A. HarrisWolfson School of Mechanical and Manufacturing Engineering, LoughboroughUniversity, Loughborough, Leicestershire, LE11 3TU, UKAbstractUltrasonic Additive Manufacturing (UAM) enables the integration of a wide variety of components into solid metal matrices due to a high degree of metal plastic flow at low matrix bulk temperatures. This phenomenon allows the fabrication of previously unobtainable novel engineered metal matrix components.The aim of this paper was to investigate the compatibility of electronic materials with UAM, thus exploring an entirely new realm of multifunctional components by integration of electrical structures within dense metal components processed in the solid-state.

Three different dielectric materials were successfully embedded into UAM fabricated metal-matrices with, research derived, optimal processing parameters.The effect of dielectric material hardness on the final metal matrix mechanical strength after UAM processing was investigated systematically via mechanical peel testing and microscopy. The research resulted in a quantification of the role of material hardness on final UAM sample mechanical performance, which is of great interest for future industrial applications.1. IntroductionUltrasonic Additive Manufacturing (UAM) utilizes ultrasonic vibrations to weld metal foils layer by layer, and then periodically applies Computer Numerical Control (CNC) machining to produce a 3D solid structure [1]. Fig. 1 shows a schematic of the UAM weld system. This weld system uses a textured sonotrode to ultrasonically oscillate and apply a compressive normal force to the metal foils to be welded.

Fig. 1 shows a schematic of the UAM weld system. This weld system uses a textured sonotrode to ultrasonically oscillate and apply a compressive normal force to the metal foils to be welded. This oscillation causes friction and deformation at the interface between the upper foil and the substrate which breaks metal surface oxide films and results in a clean, intimate surface contact. Metallurgical bonding at the interface is achieved in the solid state.Due to a high degree of metal plastic flow at a low processing temperature, UAM enables the integration of electrical components into solid metal matrices in a layer by layer fashion. Several electrical components, such as optical fibre sensor [2], pre-packaged temperature sensor [3], and direct-written circuitries [4], have previously been encased into metal matrices via UAM.

Several electrical components, such as optical fibre sensor [2], pre-packaged temperature sensor [3], and direct-written circuitries [4], have previously been encased into metal matrices via UAM. The potential for embedding freeform electrical components and circuitries within dense metal components processed by UAM presents many interesting opportunities for arange of technology sectors from aerospace to civil engineering. The successful integration of freeform circuitry and sensors into solid metal matrices would permit the manufacture of novel multifunctional complex structures that have been previously unobtainable.Fig. 1: Schematic of the Ultrasonic Additive Manufacturing (UAM) weld system.To realise these new structures a hybrid multi manufacturing systems approach was need. The materials and capabilities of the necessary manufacturing systems need to be investigated in unison to identify the key factors.

The materials and capabilities of the necessary manufacturing systems need to be investigated in unison to identify the key factors. This paper documents an investigation into the compatibility of printed electronic materials with the UAM process.2. Methodology2.1 Sample PreparationA 5 mm thick and 30 mm wide aluminium (Al) 1050 H14 plate was used as a base plate for the UAM process, and two Al 3003 H18 foils with the thickness and width of 100 µm and 25.4 mm respectively were sequentially welded onto the base plate to create the initial metal matrix (Fig. 2(a)). The UAM Alpha 2 system with parameters of 1600 N normal force, 25 µm sonotrode amplitude, and 20 mm/s welding speed, respectively, was used. These parameters were obtained via systematic tests and prior studies focused on the UAM of 3003 H18 aluminium [5-7] to produce high strength Al-Al welding.Fig.

These parameters were obtained via systematic tests and prior studies focused on the UAM of 3003 H18 aluminium [5-7] to produce high strength Al-Al welding.Fig. 2: The sample preparation method that combines screen printing electrical materials on to UAM surfaces and subsequent UAM encapsulation into solid metal matrix.Three dielectric inks, LuxPrint® 8153 from DuPont™, 520 Series Soldermask made by Technic, and Imagecure® AQ XV501T-4 of Sunchemical® were employed in the experiments and screen printed on the UAM deposited Al substrate by DEK 265 horizon printer (Fig. 2(b)). After that, all three inks were cured/solidified thermally as per the manufacturer’s instructions [8-10]. The final cure sample dimensions are 38 mm long, 3 mm wide, and around 45 µm thick (Fig. 3). Thin and small area dielectric film may be desirable for embedding via UAM process.

The final cure sample dimensions are 38 mm long, 3 mm wide, and around 45 µm thick (Fig. 3). Thin and small area dielectric film may be desirable for embedding via UAM process. However, in future work dielectrics will be used to encapsulate printed electrical conductive structures to build electrical devices, so dielectric films with the similar dimensions of finished electrical devices were fabricated and tested in this paper.Fig. 3: 520 Series dielectric film deposited on UAM processed Al substrate.The screen printed dielectric layers were then encased by ultrasonically welding another Al 3003 H18 foil on the substrate (Fig. 2(c)).

2(c)). Two combinations of control parameters with different UAM processing energy were applied to embed the electronics: a higher UAM energy combination was used to fabricate one set of metal matrices: 1600 N normal force, 25 µm sonotrode amplitude, and 20 mm/s welding speed; and a lower UAM energy set was 800 N normal force, 15 µm sonotrode amplitude, and 10 mm/s welding speed. The low energy combination was obtained by systematically embedding three dielectric films until there was no fracture caused by UAM in Al cover foils.2.2 Testing MethodsVickers Micro-HardnessThe hardness of the dielectric material was thought to be of importance for embedding as the material would need to be sufficiently hard to resist excessive deformation during the embedding process but also ductile enough to prevent cracking and disintegration when exposed to the cyclic loading of the UAM process. Hardness of dielectric materials was not found to be published or common knowledge.

Hardness of dielectric materials was not found to be published or common knowledge. In order to establish the hardness of the printed dielectric materials a Mitutoyo HM-124 micro-Vickers hardness testing machine was used to measure the hardness of the three cured dielectric materials (Fig. 4 (a)). Indentations were performed under the load control mode that increased the load to a pre-set value and held for a certain time to reduce the influence of creep (Fig. 4(b)). The maximum load applied to all three dielectrics was0.05 kg and the load duration time was set to be 10 s. The two diagonals of the indentation left in the surface of the material after removal of the load were measured using a microscope (Fig. 4(c)). The area of the sloping surface of the indentation was calculated. The Vickers hardness is the quotient obtained by dividing the kilogram-force load by the square millimetre area of indentation.Fig.

The Vickers hardness is the quotient obtained by dividing the kilogram-force load by the square millimetre area of indentation.Fig. 4: (a) The photo of Mitutoyo HM-124 micro-Vickers hardness testing machine; (b) micro- Vickers hardness test on 520 Series dielectric film; and (c) an indentation made on 520 Series dielectric film under the load of 0.05 kg.Mechanical Peel TestingPeeling tests were performed in accordance with BS EN 2243-2 2005. For each dielectric embedded with both the higher and lower UAM settings three samples were peeled and the average peel load was calculated. Samples with no embedded dielectric were also fabricated using both lower and higher parameter combinations and tested in the same way as the dielectric containing samples.

Samples with no embedded dielectric were also fabricated using both lower and higher parameter combinations and tested in the same way as the dielectric containing samples. By comparing the peeling loads for the samples with and without dielectric films any potential mechanical strength change could be clearly revealed.Optical MicroscopyEmbedded samples were cross-sectioned and investigated via optical microscopy. Linear Weld Density (LWD) was used to describe the percentage of bonded area along the weld interface, the proportion of encased dielectric area to the total embedment interface and is expressed as:(cid:1838)(cid:1849)(cid:1830)(cid:4666)%(cid:4667) (cid:3404)(cid:3013)(cid:3277) (cid:3013)(cid:3284)(cid:3400) 100 (1)where Lb and Li are the bonded area length and the bond interface length, respectively.For each dielectric embedded with both the higher and lower parameter combination, two samples were cross-sectioned and microscopically investigated.

Each sample was cut intofront middle and rear sections (Fig. 5) which were mounted in epoxy resin and then gradually grinded and polished to 0.05 µm finish. An Olympus BX60M optical microscope with a ×100 magnification lens was used to take images for LWD calculation. For each mounted section, six images of weld interface were taken from both sides of the embedded dielectric.Fig. 5: Method of sample sectioning and microscopic analysis to determine the linear weld density via optical microscopy.3. Results3.1 Mechanical Peel Testing and Vickers HardnessThe peel testing was performed as described in section 2.2. The average maximum peeling load for the three dielectric films embedded with the two combinations of UAM parameters were plotted in Fig. 6.Fig.

The average maximum peeling load for the three dielectric films embedded with the two combinations of UAM parameters were plotted in Fig. 6.Fig. 6: The average maximum peeling loads for three different hardness dielectric films embedded with two combinations of UAM parameters.For each combination of UAM process parameters, the peeling load increased with the dielectric hardness. For each dielectric, the samples encased with higher UAM process energy (1600 N force, 20 mm/s speed, and 25 µm) exhibited larger peeling load than those made by lower UAM process energy (800 N force, 10 mm/s speed, and 15 µm). Compared with the samples embedded without dielectric film that were used to characterize the mechanical strength of aluminium matrices, the peeling load of the harder dielectric such as 520 Series reduced slightly by around 8 %, whereas for the softer dielectric (8153) a dramatic reduction (more than 20 %) was observed.

Moreover, there was a significant standard deviation (± 16 N) on the peeling load of 8153 encased by lower UAM process energy.3.2 Optical Microscopy and Vickers HardnessThe LWD was measured and then calculated for each dielectric via the microscopicanalysis of sample cross-sections. The average LWDs are demonstrated graphically in Fig. 7.The general tendency was that higher UAM processing energy and a harder dielectric resulted in a higher average LWD. It is noticeable that average LWD of 8153 embedded using lower UAM processing energy also showed a large deviation (around 27 % in percentage). Both of these were in accordance with the peeling load results.Fig. 7. Linear weld density for three different hardness dielectric films embedded with two combinations of UAM parameters.3.3 DiscussionThe mechanical strength of UAM samples with embedded dielectric was enhanced following the increase of dielectric hardness (Fig.

6); this result correlates with increased LWDs, shown in Fig. 7, which also increased with dielectric hardness. This phenomenon was likely caused by the deformation of dielectric film during UAM embedding. Hard dielectric had highresistance to deformation when the UAM load was applied, so the shape of the dielectric film could be maintained. On the contrary, soft dielectric as in 8153 was found to be squashed and squeezed into the Al-Al weld area (Fig. 8). With large UAM embedding energy 8153 was deformed and squeezed uniformly along the welding direction. This resulted in a dramatic reduction of LWD in Fig. 7. Due to sufficient UAM processing energy and the regular spread of dielectric film, the standard deviations of LWD was less than ± 7 %. However, in the case of low UAM energy, a significant deviation of LWD occurred as the UAM processing energy was not sufficient to suppress the Al cover layer to adapt the irregular dielectric material expansion.

In both cases, the squash of dielectric layer caused the mechanical strength degradation of the UAM embedded structure.Fig. 8. The squashing of 8153 into the Al-Al weld area due to UAM prcessing.By observing all the samples after peel testing, no bonding was found in the interface between the dielectric films and the Al cover foils, which indicated that there was no melting of dielectric material during UAM embedding.4. ConclusionsThis work successfully deposited via screen printing and embedded via UAM a range of dielectric films thus demonstrating their compatibility with the UAM process. This result makes it clear that the direct embedding of screen printed electronic materials within solid state additively manufactured metal matrices is feasible.Peel testing and cross-sectional microscopy showed that the mechanical strength of dielectric containing UAM samples increased with the dielectric materials hardness.

The greater deformation of the softer dielectric materials caused by UAM processing was found to be the reason for the reduction in mechanical strength. For each dielectric, it was also observed that mechanical strength of the samples was improved when using higher UAM processing energy.In the near future, an electrical conductive structure will be deposited and encapsulated within the dielectric film before UAM encapsulation. This work is expected to then lead onto thecreation of UAM devices with integrated electrical components thus realising a multifunctional metal matrix composite.AcknowledgementThis work was supported by the Engineering and Physical Science Research Council(EPSRC) as part of the Centre for Innovative Manufacturing in Additive Manufacturing.References[1] White, Dawn R. "Ultrasonic consolidation of aluminum tooling." Advanced materials &processes 161, no. 1 (2003): 64-65.[2] Kong, Choon Yen, and Rupert C. Soar.

PLORING VACUUM CASTING TECHNIQUES FOR MICRON AND SUBMICRON FEATURESM. Denoual*, P. Mognol**, B. Lepioufle* *Biomis-SATIE ENS-Cachan antenne de Bretagne, Campus Ker Lann, av Robert Schumann 35170 Bruz, France **IRCCyN Nantes, FranceReviewed, accepted August 25, 2004AbstractA study of resolution limits in standard rapid prototyping vacuum cast molding processes and adaptation of this technique to reach submicron accuracy is proposed. Micro-fabrication technologies are used to fabricate micron and submicron high aspect-ratio patterns on the original parts. The molding of the original parts is optimized to allow replication of submicron features. In carefully exploring materials and surface treatments, cast parts are successfully replicated with submicron and high aspect ratio micron structures.

In carefully exploring materials and surface treatments, cast parts are successfully replicated with submicron and high aspect ratio micron structures. These encouraging results enable the use of such processes for micro- and nano-systems applications and open the door to development and production of low cost, high resolution biochips.IntroductionMicrosystems including Micro-ElectroMechanical Systems (MEMS), micro total analysis systems (µTAS), micro arrays and biochips, are now well known as an emergent field in research and industry. The micro-fabrication technologies for microsystems result mainly from microelectronics technologies based on silicon processing [1] taking advantage of high volume production of microelectronics industry equipment. However, other materials can be preferred in microsystems fabrication for specific material properties or fabrication cost. This is the case for polymers newly introduced in the microsystem field through biochemical applications.

This is the case for polymers newly introduced in the microsystem field through biochemical applications. They are particularly fitted to biochemical microsystems that require easily functionnalisable materials. Transparency is often a critical property, particularly for biomicrosystems putting silicon out of the race. Moreover, bio-chemical microsystems that serve as vehicles for performing chemical and biochemical reactions, don’t require on-board complex electronics except in some cases metallic electrodes, this eliminates silicon for those applications. And finally, the silicon micromachining techniques, such as deep-etching, are still low-volume, high-cost techniques. This need for polymer microsystems yields to the adaptation of plastic industry techniques such as hot-embossing [3-6] and injection molding [7]. The resulting techniques, micro-hot embossing and micro-injection molding, present the same characteristics as their macro equivalent.

The resulting techniques, micro-hot embossing and micro-injection molding, present the same characteristics as their macro equivalent. They require a large investment in time and money before the first part is realized and they become cost-effective only after several ten’s of thousands parts are produced.Thus, there is a niche prior to mass production for a low cost and low volume polymer microsystem part production technique suitable for rapid prototyping.In this work, the vacuum casting process was adapted to the replication of structures with accuracy reaching micron and submicron features. It has been proved that this technique can be used for the micro-fabrication of microsystems. Replicated microstructures were obtained by classic micro-fabrication the replication of microstructures have been achieved by minor improvements of the conventional vacuumtechniques. Successful results for346casting process.

Successful results for346casting process. Using the improved vacuum casting techniques for biochip fabrication has shown that this technique is suitable for pilot lot production because of its low cost and fast design to product flow.This paper aims to show the work and improvements performed in adapting the vacuum casting technique to microstructure duplication. The paper will show micro-fabrication technologies used to realize the original parts, the vacuum casting process, and the performances of vacuum casting technique for the replication of micron structures, high- aspect ratio and sub-micron structures.Materials and MethodsThe experimental step demonstrated in this paper was that : micro-structured original parts were subjected to the vacuum casting process and then duplicated parts were characterized in terms of dimensions and aspect-ratios.1.

Microstructures fabricationIn order to qualify the vacuum casting process at micron-scales, microstructures needed are obtained through microelectronics industry fabrication techniques and micro-fabrication technologies. The following are the techniques used.a. Photolithography for micron scale structures (2µm to 300µm) withmedium aspect-ratios (2:1 up to 1:6)Standard microelectronic industry photolithography using a negative-tone photoresist (Microchem SU-8) was employed to realize the original parts. This process enables the patterning of micro-scale features by crosslinking of UV-exposed regions in the SU-8 photoresist [8]. This process, derived from microelectronics technologies, is presented in Figure 1. A standard silicon wafer is used as a substrate for the fabrication process (step 1.). SU-8 photoresist is spun on the wafer (step 2.). The photoresist viscosity and the spinning rotation speed determine the resulting thickness of the layer.

SU-8 photoresist is spun on the wafer (step 2.). The photoresist viscosity and the spinning rotation speed determine the resulting thickness of the layer. In this work, two thicknesses are used: 20µm and 100µm. Once applied, the photoresist is baked to evaporate the solvents. The photoresist coating is then exposed to UV light through a chrome mask (step 3) in order to pattern the photoresist. Exposure time varies upon the photoresist thickness. Patterns designed on the chrome mask are geometrical shapes (rectangles and circles) with dimensions ranging from 2µm to 300µm. The UV-exposed photoresist is reticulated through a second bake. The non-exposed zones are dissolved in developer (step 4). Figure 2 shows the typical negative taper profile of an SU8 photoresist pattern on a 100µm high 20µm diameter tower, intentionally a broken tower is showed to facilitate the observation.

Figure 2 shows the typical negative taper profile of an SU8 photoresist pattern on a 100µm high 20µm diameter tower, intentionally a broken tower is showed to facilitate the observation. This observation underlines the fabrication limitations of the negative photoresist process that enables high aspect ratio structures of up to 1:5 or 1:6.347Figure 1 : Negative photoresist process.Figure 2 : Negative photoresist towers.b. Oxide layer patterning for sub-micron features (down to 300nm)The process used for the realization of the sub-micron features is a standard microelectronics photolithography and patterning of an oxide layer traditionally used as an insulation layer in semiconductor process.

Lithography offers in-plane sub-micron precision on dimensional scales from micron to millimeter.Figure 3 : oxide layer patterning process.Figure 4 : oxyde dots (1µm in diameter, 300nm thick).The figure 3 describes the process used to obtain the high-aspect ratio structure used for duplication. A silicon wafer is first oxidized (step 2). Then a standard microelectronics photolithography process is performed. Positive-tone photoresist is spin-coated on the wafer (step 3) and exposed to UV light through a chromium mask (step 4). Exposed photoresist is then developed leaving the unexposed regions (step 5). The remaining photoresist acts as a mask during the etching of the oxide layer in acid bath or in plasma of a reactive ion etching (RIE) machine (step 6). The thickness of the oxide layer is 300nm. Photoresist is stripped (step 7). The figure 4 shows matrix of oxide dots with 1µm in diameter and 300nm in height.c.

The thickness of the oxide layer is 300nm. Photoresist is stripped (step 7). The figure 4 shows matrix of oxide dots with 1µm in diameter and 300nm in height.c. Silicon deep etching for high aspect-ratio structures (1:50)Silicon deep-etching was introduced at the beginning of the 90’s as a new technological process for micromachining of silicon. Specific tools were developed to perform this process called deep reactive ion etching (deep-RIE) because it is an adaptation of the RIE semiconductor industry process to the etching of deep structures. Since that time, deep-348etching has taken an increasing role in MEMS processing especially for actuators and sensors [9].

Since that time, deep-348etching has taken an increasing role in MEMS processing especially for actuators and sensors [9]. Deep-RIE has not yet entered high-volume manufacturing because of single wafer processing and wafer non-uniformity of the etching reaction.Figure 5 : silicon deep-etching Bosch process.Figure 6 : SEM image of a polished microsection of a deeply etch silicon wafer.Figure 5 describes the process used to obtain the high-aspect ratio structure used for duplication. An oxide layer acts as a mask during the deep-etching of the silicon. The oxide layer patterns are obtained through the process described in the previous paragraph. The deep- etching process developed by Bosch takes advantages of a side effect of glow discharges, the tendency to create polymeric species by chemical cross-linking. In fact, most plasma processes are a critical race between deposition of polymeric material from the plasma and the removal of material from the surface.

In fact, most plasma processes are a critical race between deposition of polymeric material from the plasma and the removal of material from the surface. In well-designed plasma chemistries, removal dominates. However, in the process developed for deep-etching, the deposition of polymer from the etchant is used to a great advantage. The etch proceeds in alternating steps of reactive-ion etching in an SF6 plasma (step 3) and polymer deposition from a C4F8 plasma (step 2). During the etch step process, the polymer is rapidly removed from the bottom of the feature but lingers on the sidewall, protecting it from the SF6 etchant. The etching depth is controlled thanks to process duration time, i.e. the number of etching cycles (repetition of step 2 and 3).Figure 6 shows an example of deeply etched silicon wafer. We can observe that the depth of the trenches depends on the opening of the trench. High aspect ratio (up to 1:50) can be performed with this technique.

We can observe that the depth of the trenches depends on the opening of the trench. High aspect ratio (up to 1:50) can be performed with this technique. The deep-etching technique tends to form structure with a slightly negative taper profile with angle between 85° and 89.5° which increase the mechanical constraints during the vacuum casting process fir the removal of the duplicated parts. Figures 7 and 8 show close-up views of the trenches that illustrate the dimensions of the walls and the scalloping effect due to the process steps.349Figure 7 : SEM image of vertical thin walls with here a 3µm wide and 84µm high wall (aspect ratio 1:28).Figure 8 : high magnification SEM image of vertical pillars.2. Silicon processing limitsMicroelectronics industry techniques facilitate the realization of micro-structured devices, but they are limited by three factors. (cid:131) The choice of material is limited to silicon-based materials.

(cid:131) The choice of material is limited to silicon-based materials. This is a real limitation when considering biological or chemical applications because of transparency and biocompatibility requirements. In addition, silicon is more expensive than polymers when bulky parts are concerned. (cid:131) The techniques allowing high aspect-ratios and volume structures patterning such as deep RIE techniques are still very expensive. (cid:131) Those techniques are not available for mass production because of the expensive equipment involved treating only a single wafer at a time and the processing time is long (with etching speed ~1µm/min).3. Duplication processThe vacuum casting process consists in the realization of a silicon female mould with an original part followed by resin casting under vacuum to obtain the duplicated parts. In the first experiments, the standard vacuum casting process was applied. The direct application of this process is illustrated in Figure 9.

In the first experiments, the standard vacuum casting process was applied. The direct application of this process is illustrated in Figure 9. A Sprue bushing channel and vent are joined to the original part to be duplicated (step 1.). A female mould is made of silicone around the original part (step 2.). VTV 750W silicone form MCP is first mixed 2 minutes in a ratio of 10:1 with the hardener solution (CAT 750) and then degassed 15 minutes prior to pouring over the original part. Curing time is 6 h at 70°C. After the bake, the silicone female mould is opened and the original part withdrawn (step 3.). Female mould is then ready for casting. MCP 6091 resin was used for casting under vacuum in the female mould. First MCP 6091 part B is degassed under vacuum for 10 minutes. Then MCP 6091 part A and MCP 6091 part B are mixed in a ratio of 1:1.8 (A:B) during 1 minute. The mixed solution is then cast in the female silicone mould (step 4.).

Then MCP 6091 part A and MCP 6091 part B are mixed in a ratio of 1:1.8 (A:B) during 1 minute. The mixed solution is then cast in the female silicone mould (step 4.). After 2 h bake at 70°C, the resin is hardened and the resin duplicated part is withdrawn from the mould (step 5.).350Figure 9 : Vacuum casting procedureFigure 10 : improved processIt appeared that the direct application of the vacuum casting process was not good enough for microsystem applications as illustrated below. The poor surface roughness was due to sweating of the silicone used for the female mould fabrication. Hence, the first improvement was to use a high quality silicone to realize the female mould. PDMS (PolyDiMethylSiloxane, Sylgard RTV184 from Down Corning) was chosen. This silicone is often used in micro- fabrication technologies for micro-molding processes [10]. The PDMS curing process is similar to the VTV 750W.

This silicone is often used in micro- fabrication technologies for micro-molding processes [10]. The PDMS curing process is similar to the VTV 750W. The PDMS prepolymer is mixed in a ratio 10 : 1 with the hardener solution and then degassed 15 minutes in vacuum prior to pouring. Curing time is quicker with 1 h at 70°C.Another improvement to prevent rupture of the silicon female mould during the withdrawal of the original part, in case of high aspect ratio structure or slightly negative taper profiles, is the deposition of an anti-adhesive coating on the surface of the original part. Indeed, a “teflon”- like layer can be deposited in CHF3 plasma by a standard reactive ion etching (RIE) machine. This treatment is particularly important for the reproduction of high aspect-ratio structures. The improved process is presented in Figure 10.Results and performances1.

This treatment is particularly important for the reproduction of high aspect-ratio structures. The improved process is presented in Figure 10.Results and performances1. Micron features, medium aspect ratioFigures 11.b,d and f illustrate the results obtained when the classic macro vacuum casting process is applied to micro-scale structures. The replication is not correct. Dimensions are not reproduced and shapes are altered. In addition the resin has a poor surface quality. Surface quality can be compared to the original part surface quality in Figure 11.a,c and e. Original parts have a flat surface whereas duplicated parts have variegated surfaces, and straight lines on the original part become wavy on the duplicated part. Those preliminary results were not351good enough to consider the use of vacuum casting for microsystem applications.(a)(c)(f) Figure 11 : The pictures on the right correspond to the original micron scale structures.

On the left, the pictures show the duplication results. SU8 cross with 300µm wide and 100µm high walls (a), resin replication of the cross (b), SU8 original towers (e), duplicated towers (f, scale bar is 100µm). The patterns are not accurately duplicated here and the surface quality is poor.(e)The improvements presented in the previous paragraph were applied (anti adhesive coating and high quality elastomer for the mould) and Figures 12.a,b,c and d show the results obtained. The surface quality is comparable to the surface of original part. Moreover, the structures are accurately duplicated.352(b)(d)(a)(c)Figure 12 : SEM image of duplicated structures. Field of towers (a), set of walls (b), negative taper profile towers (c) with 10µm diameter, 20µm high, and a 11.6µm wide, 20µm high duplicated wall (d).Using this technique, features ranging from 10 to 300 µm were successfully replicated. Aspect ratio of the realized structures varied up to 1:6.

Aspect ratio of the realized structures varied up to 1:6. The smallest structures of the original parts were successfully duplicated and the duplicated dimensions were limited only by the dimensions of the original parts. Duplication error is estimated to be about 1%.The slightly negative taper profile were also successfully replicated thanks to the elasticity of the mould. Such profiles are not allowed with hot embossing or injection moulding since they lead to structural breakdown of the mould or prevent part withdrawal.2. Submicron featuresTo qualify the replication of submicronic features, the improved vacuum casting process has been performed for the duplication of oxide layers patterned on silicon. Figure 13 presents original oxide and silicon parts against their resin vacuum casting replication.

Figure 13 presents original oxide and silicon parts against their resin vacuum casting replication. The smallest pattern dimensions are 0.8µm and the oxide layer thickness is 300nm.353(b)(d)(a)(d) Figure 13 : SEM images of original oxyde and silicon structures and their resin duplication. Geometrical shapes patterned in oxide (a), resin replicated shape (b), small oxide dot (c) and its replication in resin (d).(c)The results demonstrate that submicron structures can be replicated using the vacuum casting improved process. Features down to 800nm wide and 300nm high have been successfully replication using the vacuum casting improved process. The limitation here is due to the process environment. The vacuum casting process is operated in ambient environment without any particular dust class control. Therefore some dust is visible on the surface of the duplicated structures. For nano-replication using vacuum casting technique, clean-room environment is required.3.

Therefore some dust is visible on the surface of the duplicated structures. For nano-replication using vacuum casting technique, clean-room environment is required.3. High aspect ratioIn order to characterize the performances of the vacuum casting process, deep-etching of silicon was used to realize high-aspect ratio (up to 1:50) silicon patterns. The structures utilized for replication were walls, trenches, cylinder and cubes, with dimensions down to 2µm wide and walls, up to 100µm in height. The results are presented in the figure 14.354(b)(a)(c)(f) Figure 14 : SEM images of high-aspect ratio silicon structures (a,c and e) and their resin replication through vacuum casting (b,d and f). A 6µm wide and 100µm high wall (b), 20µm diameter and 100µm towers (d).(e)Visualization of polymer structures with a SEM requires the coating of the sample surface with a thin conductive layer (chromium-gold).

However a side effect of the high aspect-ratio structures is creation of shadows during the deposition of this layer, which causes brightness variations on the images of resin structures compared to the images of silicon structures.The maximum aspect ratio obtained in a replicated structure is 1:16. One could probably improve this aspect ratio with positive profile high aspect ratio structures.Structures with an aspect ratio of 1:20 have been replicated as illustrated in the figure 15. Indeed the replicated cylinders are 5µm in diameter and 100µm high. But the material stiffness seems insufficient to maintain the shape. This implies future work on the material choice and processing to improve aspect-ratio.355(b)(d)(b) Figure 15 : SEM image of resin duplicated patterns. 5µm diameter and 100µm high cylinders.

This implies future work on the material choice and processing to improve aspect-ratio.355(b)(d)(b) Figure 15 : SEM image of resin duplicated patterns. 5µm diameter and 100µm high cylinders. Low magnification (a), high magnification close-up view (b).(a)ConclusionThe adaptation to the replication of microstructures of a conventional rapid prototyping technique was investigated. According to the experimental results, a conventional vacuum casting process with minor improvements can be applied to the fabrication of microstructures. High quality materials are essential to overcome sweating and adhesion problems found at the micron scale. The duplications of classical test structures showed sub-micron (down to 300nm) and high aspect-ratio (up to 1:16) performances. The improved vacuum casting process allows accurate replication of microstructures made with standard micro- technologies. A cost study has been performed for a specific bio-chemical microsystem application.

A cost study has been performed for a specific bio-chemical microsystem application. The results of the study, illustrated in the figure 16, stress the cost-effective domain of this technique : from a few parts to a few thousand parts. Therefore, improved vacuum casting meets the need for small series of micro-structured systems made of polymers. Because of its high speed to first part and low-cost, the adaptation of the vacuum casting process to micro- fabrication will lead to a promising technology for MEMS application. This is indeed a suitable technique for rapid prototyping of polymer microsystems dedicated to biology or chemistry when compared to hot embossing or micro-injection molding techniques that require significant more time and money.Figure 16 : part cost versus production volume compared for injection molding, micro- fabrication technology and vacuum casting process.356REFERENCES [1] J.W. Gardner, V.K. Varadan, O.O.

PLORING VARIABILITY IN MATERIAL PROPERTIES OF MULTI-MATERIAL JETTING PARTSLindsey B. Bass, Nicholas A. Meisel, and Christopher B. Williams Design, Research, and Education for Additive Manufacturing Systems Laboratory Department of Mechanical Engineering, Virginia TechABSTRACTWith Additive Manufacturing (AM) capabilities rapidly expanding in industrial applications, there exists a need to quantify materials' mechanical properties to ensure reliable performance that is robust to variations in environment and build orientation. While prior research has examined process-parameter and environmental effects for AM processes such as extrusion, vat photopolymerization, and powder bed fusion, existing similar research on the material jetting process is limited.

Focusing on polypropylene-like (VeroWhitePlus) and elastomer-like (TangoBlackPlus) materials, the authors first characterize the anisotropic properties of six different gradients produced from mixing the two materials in preset quantities. Three build orientations were used to fabricate parts and analyze tensile stress, modulus of elasticity, and elongation at break for each material. The authors also present results from an investigation of how aging of parts in different lighting conditions affects material properties. The results from these experiments provide an enhanced understanding of the material behaviors relating to material jetting process parameters and can inform material selection when manufacturing load- bearing parts.1. INTRODUCTIONAdditive Manufacturing (AM) is quickly evolving from a method for prototyping to a desired alternative for manufacturing end-use products.

INTRODUCTIONAdditive Manufacturing (AM) is quickly evolving from a method for prototyping to a desired alternative for manufacturing end-use products. Because of the nature of a layer-by-layer fabrication process, customizable artifacts are achievable that save material, time, and cost compared to traditionally-manufactured parts. In particular, the material jetting process works by selectively depositing droplets of build material to form parts [1]. An example is the PolyJet material jetting system commercialized by Stratasys [2] that is capable of simultaneous deposition of multiple photopolymer resins. In the PolyJet process, a print block consisting of the inkjet heads deposits the support or build materials in drop-by-drop deposition patterns, which are smoothed by a roller and cured by a UV lamp (Figure 1).Figure 1.

Representation of the PolyJet Printing ProcessWhile material jetting processes, such as the PolyJet process, offer unique multi-material capabilities, research is still needed to identify the impact that process variations have on the quality of the final manufactured part. For example, this style of layer deposition can introduce build defects and inconsistencies due to how the material droplets are spread and bond. In order to ensure that load-bearing, end-use, material-jetted products meet the required specifications (which could include having high ultimate tensile stress, directionally-independent ultimate tensile stress, or strength longevity), variations in the material properties must be examined.In this paper, two common materials for the Stratasys Objet350 Connex were studied: a polypropylene-like material, VeroWhitePlus (VW+), and an elastomer-like material, TangoBlackPlus (TB+), as well as gradients produced from drop-by-drop dithering patterns of the materials.

To identify variations in these digital materials, mechanical properties were investigated with respect to build orientation and part age, two key factors that have the potential to significantly affect tensile stress, elastic modulus, and elongation at break.Related research has been divided into three parts. Section 1.1 outlines the main contributors to orientation effects in various AM technologies. Since each fabrication process is different, material properties are also expected to differ. In Section 1.2, advances in the material jetting process are noted, and these findings guided the direction of this experiment.

In Section 1.2, advances in the material jetting process are noted, and these findings guided the direction of this experiment. Section 1.3 transitions to aging effects, which is currently limited in literature for AM photopolymer materials; however, the vat photopolymerization findings are important to consider and should be compared to the material behaviors witnessed in this material jetting study.1.1 Anisotropy in AMExisting literature has explored the effects of orientation on mechanical properties for multiple AM processes. Kotlinski offers a review comparing AM materials with an extensive list of mechanical properties [3]. Among powder bed fusion, material extrusion, binder jetting, vat photopolymerization, sheet lamination, and material jetting technologies, anisotropic effects exist, but are shown to differ.

Among powder bed fusion, material extrusion, binder jetting, vat photopolymerization, sheet lamination, and material jetting technologies, anisotropic effects exist, but are shown to differ. Puebla and coauthors investigated orientation with vat photopolymerization and found that the specimens built flat on the build plane had significantly lower ultimate tensile stresses and elastic modulus values compared to parts built vertically or on an edge [4]. Lee and coauthors looked at multiple build orientations in the extrusion, binder jetting, and nano composite deposition system processes [5]. Binder jetting was the only method producing low axial compressive strength, and extrusion had the highest axial compressive strength. The compressive tests further confirmed that build direction is a significant parameter affecting material properties. Zeleny and coauthors compared extrusion and material jetting with ABS plastic and an ABS-like material respectively [6].

Zeleny and coauthors compared extrusion and material jetting with ABS plastic and an ABS-like material respectively [6]. Extrusion parts printed on their edges had higher tensile strengths than parts printed flat. Among the three orientations for the material jetting parts (flat parts printed along and perpendicular to the build direction and parts printed on their edge along the build direction), the tensile strengths more than doubled compared to extrusion parts. Flat parts printed perpendicular to the build direction had the highest tensile strength, while parts printed on an edge had the lowest.1.2 Orientation Effects in Material JettingConcerning studies solely focusing on the material jetting process, Pilipović and coauthors thoroughly explored a variety of mechanical properties of VeroBlack, VeroBlue, and FullCure720 materials and found FullCure 720 to perform the best in maximal flexural strength [7].

Similarly, Singh and Singh compared three jetted materials and were able to determine that VeroWhite material fabricated horizontally compared to at 45- or 90-degree angles with the build bed was the most dimensionally-accurate and cost-effective material as opposed to VeroBlue and Fullcure 720 [8]. Blanco and coauthors also directly explored slant angles with strips printed in seven acute-angle orientations [9]. The relaxing modulus was highest at 0 and 90 degrees, and their trends revealed the potential for a shielding effect due to how the support material is UV-cured. Barclift and Williams used a design of experiments to test factors of XY- orientation (in the build plane) and Z-orientation (height direction) [10]. While they were unable to establish statistically significant trends in tensile strength and tensile modulus pertaining to orientation, their work highlights how Z-orientation can affect the dimensional accuracy of a part produced via material jetting.

Adamczak and coauthors analyzed three orientations of VeroWhite material and discovered anisotropy although averaged ultimate tensile stress values were similar across orientations [11]. Vertical pieces, however, were found to be the most brittle. Using FullCure 720, Kesy and Kotlinski also witnessed an anisotropic trend but with different findings: parts oriented with the longest dimension along the build direction and second longest along the Z-axis were the strongest parts, parts oriented vertically were next strongest, and the weakest were parts with the longest dimension along the build direction and the second longest perpendicular to the build direction in the build plane [12].Orientation results in prior literature clearly differ by AM process: while flat parts had lowest ultimate tensile stresses in vat photopolymerization, only edge-oriented extrusion parts experienced significantly higher ultimate tensile stresses.

Material jetting research has often produced inconsistent anisotropic trends, but some tests have shown parts on their edge being strongest. Besides considering the build process differences, different materials do not maintain similar trends, which further emphasizes the difficulty of material characterization and the necessity for continued research. In this study, both VW+ and TB+ will be examined to learn about the behavior of rigid and flexible materials in multiple orientations and how (if at all) the mechanical properties vary. To the authors’ knowledge, this is the first orientation analysis of digital materials involving TB+.1.3 Aging in Photopolymer AM ProcessesAside from orientation, another important factor potentially affecting the strength of end-use AM parts is age. The vast majority of prior research into aging of AM parts is with the vat photopolymerization process.

The vast majority of prior research into aging of AM parts is with the vat photopolymerization process. Tröger and coauthors investigated aging of multiple acrylate- based resins to more clearly understand the materials’ behaviors as they apply to engineering and biological applications [13]. By considering thermal aging, humidity aging, and UV aging separately, the authors were able to determine that while aging mainly depended on the material, humidity caused swelling and a reduction in mechanical properties. Furthermore, higher temperatures and short wavelength light irradiation affected material color and mechanical properties. Ottemer and Colton examined epoxy-based resins in a seven-week study with four different relative humidity scenarios [14]. While aging did not yield significant trends in mechanical properties, specimens in the more humid environments noticeably absorbed water and deteriorated in strength.

While aging did not yield significant trends in mechanical properties, specimens in the more humid environments noticeably absorbed water and deteriorated in strength. Using a 24-day aging cycle, Mansour and coauthors also used an epoxy resin and determined that as aging proceeds, the mechanical properties reach equilibrium after an improvement of tensile modulus, maximum tensile stress, flexural modulus, and strength[15]. The percent elongation at break and impact strength were shown to decrease over time. Puebla and coauthors also took a look at aging, and they used 4-, 30-, and 120-day intervals for aging experiments. Results showed that the shortest time interval corresponded to the lowest ultimate tensile stresses. Humidity significantly reduced mechanical properties as did the longest aging cycle for one of three materials investigated [4].The results from current literature fail to reach an agreement regarding one trend of mechanical property behavior over time.

To the authors’ knowledge, there is no research in aging behavior for the material jetting process. The authors have therefore chosen to focus on this process with VW+ material to make a valuable contribution to expanding knowledge on the evolution of mechanical properties with age.1.4 ContextIn this paper, the effects of orientation and aging are analyzed to determine the effect of material jetting process parameters on the mechanical properties of end-use parts. Three build orientations are explored among six photopolymer materials. In a separate investigation, one material is studied over 10 weeks, a length of time the authors felt could allow parts to set in order to yield changes in material behavior after many days. The analysis of the mechanical properties will provide knowledge about material characterization leading to a better understanding of the reliability of material-jetted end-use parts.

The analysis of the mechanical properties will provide knowledge about material characterization leading to a better understanding of the reliability of material-jetted end-use parts. Section 2 discusses the effect of build orientation on mechanical properties, and the effect of aging is analyzed in Section 3. Both of these sections begin with an outline of the experimental methods (Sections 2.1 and 3.1 respectively.) Sections 2.2 and 3.2 present the results of ultimate tensile stress, modulus of elasticity, and elongation at break. A summary of the key findings follows in Sections 2.3 and 3.3. Finally, the closure evaluates the achievements of this investigation as well as future considerations (Section 4).2. EFFECT OF ORIENTATION 2.1 Experimental MethodsTo better understand the material behavior of multi-material jetting parts, six different materials were investigated, each with different concentrations of VW+ and TB+ based on pre- arranged drop-by-drop dithering patterns.

The materials are displayed in Table 1 along with abbreviations that will be used in this paper. VW+ is a rigid, polypropylene-like photopolymer material, while TB+ is a flexible, elastomer-like photopolymer material. The flexibility of the intermediate gradient materials is directly proportional to the percentage of TB+ material present in the printed composite. The experimentation in the following section is intended to explore variability in mechanical properties of six digital materials when they are built in different orientations.Table 1. Range of Materials for the Study [16]Material (Rigid) Abbreviation VeroWhitePlus VeroGrey35 VeroGrey50VW+ VG35 VG50Material (Flexible)AbbreviationTangoBlackPlus_Shore85 TBS85 TangoBlackPlus_Shore60 TBS60 TangoBlackPlusTB+Using the ASTM D638 Type IV tensile specimen [17], mechanical properties of each digital material were evaluated for three different orientations.

Three specimens were printed for each material and orientation under investigation. Figure 2a provides the configuration of parts oriented with their longest direction along the X-axis and their second-longest direction along the Y-axis, which will be referred to as XY-parts to conform to ASTM standards [18]. Parts with the longest dimension along the Y-axis and second-longest along the X-axis (YX-parts) are shown in Figure 2b, and Figure 2c presents the ZX-parts: parts with the longest dimension along the Z-axis and the second-longest along the X-axis. These three orientations were selected out of the six possible orientations to be representative of how each of the longest directions affects mechanical properties. Additionally, many sources ([4], [6], [9], [11], [12]) set up their samples with all three orientations on the same build tray, but this study separates orientation by tray to ensure there are uniform UV-curing effects across all specimens.

When parts of different heights are placed on the same build tray, the UV irradiation has potential to over-cure parts adjacent to the current printing path [10], which could be the case on a tray with multiple orientations when only the vertical parts are still printing and while the top surfaces of the flat parts remain exposed for the remainder of the build.ZZZYXYXYX(a)(b) (c)Figure 2. Parts oriented on the build tray with various grayscales representing the different materials: a) XY-parts, b) YX-parts, and c) ZX-parts.All specimens were arranged on the build tray using the automatic placement feature with locked orientations. Each tray had three specimens of each of the six materials, and materials were randomly assigned to their placement.

Each tray had three specimens of each of the six materials, and materials were randomly assigned to their placement. In order to prevent the machine’s vibrations or the print roller from causing undesired motion in the ZX-parts during printing, a roof-like structure consisting of a thin sheet of VW+ material was designed to be printed above the specimens. This forced the machine to fully support the specimens by fabricating the parts all within a solid block of support material. Without incorporating this strategy, the ZX-parts wobbled during the build and exhibited severe print defects in their upper regions.Parts were printed in Digital Materials Mode (layer thickness of 32 microns) with matte finish. Following printing, all support material was removed via a high-pressured waterjet cleaning station. Surrounding conditions were maintained at an average of 72˚F and 27% humidity.

Following printing, all support material was removed via a high-pressured waterjet cleaning station. Surrounding conditions were maintained at an average of 72˚F and 27% humidity. Environmental conditions were maintained relatively consistent because of the potential for altered results from introducing temperature and humidity as additional variables. This has been observed with past experiments in humid environments where part strength was greatly diminished ([4], [13], [14]). To reduce potentially harmful effects from extended environmental exposure, parts were cleaned and tested immediately upon print completion.The print block for the material jetting process, which is referenced in Figure 1, moves along the X-axis from left to right. The print pass rows are defined by the white lines parallel to the X- axis on the build trays in Figure 2.

The print pass rows are defined by the white lines parallel to the X- axis on the build trays in Figure 2. The authors hypothesized that this directional print process unique to the material jetting technology would cause material properties to be dependent on orientation. XY-parts were expected to be strongest since they are jetted along the direction of tensile strain. ZX-parts were predicted to be weakest since each deposited layer has small cross- sectional areas, creating the potential for delamination. Since the XY-plane has the lowest printing resolution, poor dimensional accuracy from the discretized nature of the jetted layer fabrication could also misalign layers along the Z-axis leading to premature failure [10].Tensile testing was performed using the Instron 5984 Mechanical Testing Machine using 10kN pneumatic grips on a 10kN load cell. The pull rate was 5 mm/min.

The pull rate was 5 mm/min. Smooth jaw faced grips were used for the flexible specimens, and serrated jaw faced grips were used for the rigid specimens. Gauge length, width, and thickness of each sample were recorded in millimeters, and every 0.1 second the extension (mm), tensile stress (MPa), load (N), and tensile strain (mm/mm) were recorded.2.2 Results and DiscussionUltimate tensile stress values are reported by orientation in Figure 3. Each of the materials corresponds to the listing in Table 1 from the most rigid (VW+) to the most flexible materials (TB+). The rigid and flexible materials are displayed in separate graphs due to large differences in Y-axis scaling. Error bars indicate sample standard deviations of each data point. Similarly, Figure 4 displays the elastic modulus for each material calculated using linear regression.

Error bars indicate sample standard deviations of each data point. Similarly, Figure 4 displays the elastic modulus for each material calculated using linear regression. The elastic region of the stress-strain curve was defined where the linear regression R2 value of the curve was at least 0.97, with the exception of the TB+ ZX-parts, which had R2 values of at least 0.90.603.5) a P M( h t g n e r t S e504030) a P M( h t g n e r t S e3.02.52.0l i s n e T e t a mi t l U2010l i s n e T e t a mi t l U1.51.00.500.0VW+VG35 Rigid Material TypeVG50TBS85TBS60 Flexible Material TypeFigure 3.

While modulus of elasticity values for TB+ are similar across orientations, there is clear anisotropy among the other materials. It is possible that this anisotropy manifests itself in part due to the pre-designed dithering patterns used to create the intermediate materials. As the concentration of TB+ in the intermediate materials changes, so may the three-dimensional dithering pattern used to create the material, which may in turn cause the observed increases in anisotropy. While the effects of dithering on the mechanical properties of material jetting parts have been investigated before [19], additional research is needed to confirm this hypothesis.Although the flexible materials’ elastic modulus values parallel the trend for the flexible materials’ ultimate stresses, the rigid materials indicate that the ZX-parts have the highest modulus of elasticity with XY-parts having the lowest.

Observing the elongation at break of each part in Figure 5 can help justify this inverted trend.70)m m60(k a e r B t a n o i t a g n o E5040302010XYYXZXl0VW+VG35VG50 TBS85 Material TypeTBS60TB+Figure 5. Part elongation at break for each orientation and material (most rigid to most flexible)XYYXZXTB+Since a more elastic part has a lower modulus of elasticity, it is understandable that the rigid materials have XY-parts elongating the most while having the lowest modulus of elasticity values. Rigid XY-parts tend to also be slightly stronger likely due to how the material was jetted along the direction of tensile strain. In contrast, the rigid ZX-parts elongate the least and are the weakest, potentially because of the tendency for delamination since cross-sectional layers have small surface areas. The rigid ZX-parts have a higher modulus of elasticity verifying that they are less elastic.

The rigid ZX-parts have a higher modulus of elasticity verifying that they are less elastic. For the flexible materials, the ZX-parts elongate the most and generally have lower modulus values. The TB+ parts specifically do not vary significantly in ultimate stress nor modulus, which can most likely be attributed to the nature of the rubber-like TB+ material allowing it to exhibit isotropic behavior during fabrication. The horizontal trends in Figure 5 for the rigid materials clearly indicate that small concentrations of TB+ material do not change how much the specimens elongate. On the other hand, the more flexible parts elongate relative to their TB+ material concentration.T-tests were used to compare the material properties among the various orientations within each individual material. Shown in Table 2 are the p-values, the likelihood of the data to be recreated in a random environment. Any p-value of 0.05 or less has been darkened to signify statistical significance.Table 2.

The TBS85 differences between XY-parts and YX-parts as well as the rigid materials are not statistically significant. Additionally, the p-value of 0.23 for X vs Z in TB+ hints at an isotropic property. The lack of statistical significance in ultimate stress and elastic modulus for the rigid materialspoints to their isotropic tendency. Elongation data was also statistically analyzed, and all of the rigid materials experience significant differences in extension. For the flexible materials, there is only statistical significance when comparing XY-parts and ZX-parts.The results from the experiment were compared against the values Stratasys provides in the Digital Materials Data Sheet [16] as shown in Table 3. The elastic modulus values measured in this study are considerably lower than what is reported in the provided data sheet. There is no modulus data presented for the flexible materials.

The elastic modulus values measured in this study are considerably lower than what is reported in the provided data sheet. There is no modulus data presented for the flexible materials. Ultimate tensile stress values match up for the rigid materials, but the flexible materials’ measured values were much lower than reported values. The elongation at break data for rigid materials match well, and the flexible materials elongate much more than predicted.Table 3.

Parts oriented with the longest direction along the Z-axis tend to be the weakest. For rigid materials, parts oriented along the X-axis have the lowest moduli of elasticity and have the highest elongations at break. Flexible parts in the same direction have the highest moduli of elasticity and the lowest elongations. While the variations in flexibility in the rigid materials do not affect their elongations at break, the flexible materials elongate more with increased TB+ concentration for all orientations. Although there is a trend of statistical significance for modulus of elasticity and ultimate stress values for most flexible materials, results are inconclusive for rigid materials. For elongations, while all materials exhibit statistical significance between XY- and ZX- parts, only rigid materials have all directions significantly different.3.

For elongations, while all materials exhibit statistical significance between XY- and ZX- parts, only rigid materials have all directions significantly different.3. EFFECT OF AGING 3.1 Experimental MethodsKnowing the effect of time on the material properties of end-use parts is critical when considering the expectation of prolonged part quality. To explore the effect of time on degradation of cured photopolymer, identical part trays were fabricated in the XY-orientation, and parts were tested weekly over ten weeks. Each tray had four VW+ specimens thatunderwent automatic placement and had matte finish. Two specimens were placed on a shelf with exposure to frequent fluorescent light, and the other two parts were placed in a desiccant box with no light exposure. For each week, two specimens from each location were tested. Tensile testing parameters are identical to those outlined in Section 2, and likewise, humidity and temperature were relatively consistent.

Tensile testing parameters are identical to those outlined in Section 2, and likewise, humidity and temperature were relatively consistent. Temperatures and relative humidity of both the surroundings and desiccant box were approximately 72˚F and 27% respectively. The experimentation in the following section is intended to explore variability in the mechanical properties of VW+ as it ages to understand how the material jetting process affects part quality.3.2 Results and DiscussionThe trends in ultimate stress and modulus of elasticity for each storage location are shown in Figure 6.

“Week 0” parts were tested immediately after being printed and are considered the control specimens, and all four samples were averaged together for analysis.651050) a P M(s s e r t S el i s n e T e t a mi t l U60555045) a P M(y t i c i t s a E f o s u u d o Mll1000950900850800750Shelf40012345 Week678910700012345 Week67Desiccant Box8910(a)(b)Figure 6. (a) Ultimate stress and (b) modulus of elasticity values for VW+ over 10 weeksThe ultimate stress values depict a similar trend between the shelf and desiccant box parts. While there is a decrease in strength at the third week, the general trend increases in ultimate stress.

While there is a decrease in strength at the third week, the general trend increases in ultimate stress. It is noticed in Figure 6b that, while slightly increasing over time, the modulus of elasticity of both groups of specimens were fairly consistent over time though the elastic modulus of the desiccant box parts seems to remain slightly lower than that of the shelf parts. To quantify the difference between results of the shelf specimens versus the desiccant-box specimens, a t-test was conducted with results displayed in Table 4. With p-values of 0.05 or less indicating statistical significance, clearly, there was not a significant change in material properties between lighting conditions even though the shelf parts had potential for extra UV exposure from everyday lighting. Since the placement proved irrelevant, the data from Figure 6 was averaged together and replotted in Figure 7. Table 5 shows computed p-values comparing for significance of each week against the control.

T-test comparing each week to the control weekWeek0 vs 1 0 vs 2 0 vs 3 0 vs 4 0 vs 5 0 vs 6 0 vs 7 0 vs 8 0 vs 9 0 vs 10Ultimate Tensile Stress 0.03 0.06 0.04 0.92 0.54 0.00 0.00 0.00 0.00 0.00Modulus of Elasticity0.01 0.15 0.01 0.17 0.02 0.79 0.09 0.50 0.16 0.23Elongation at Break0.34 0.08 0.03 0.06 0.01 0.00 0.01 0.00 0.00 0.0078910Elongation at break results are presented in Figure 8 with p-values shown in Table 5. A pronounced downward trend, which is statistically significant after three weeks, points to parts becoming less elastic over time. The slight increase in modulus of elasticity values verify this trend. And just like in Section 2, these trends correspond to higher ultimate stresses. This indicates that as the material sits over time, it may be hardening, which makes it stronger but more brittle.16)m m14(k a e r B t a n o i t a g n o E e g a r e v Al121086420024 Week6810Figure 8.

Part elongation at break for VW+ over 10 weeks3.3 Summary of Key FindingsBy testing the mechanical properties of VW+ over 10 weeks, the authors are able toconclude: The material properties evaluated in this study for VW+ specimens are not significantlyaffected by different lighting conditions. Ultimate tensile stress of VW+ parts increases as parts age. Results are statisticallysignificant after five weeks. Modulus of elasticity values were not shown to change at a statistically significant level.  Elongation at break shows a significant decreasing trend over time.4. CLOSUREIn order for the material jetting process to make effective end-use parts, there must first be an understanding of how the process affects part quality. The goal of this study was to identify variations in PolyJet digital materials by looking at effects of part orientation and aging.

The goal of this study was to identify variations in PolyJet digital materials by looking at effects of part orientation and aging. This paper has explored three key mechanical properties (ultimate tensile stress, modulus of elasticity, and elongation at break) of several materials made from two common material jetting photopolymers: polypropylene-like VeroWhitePlus (VW+) and elastomer-like TangoBlackPlus (TB+). Looking at orientation, XY-parts of all materials tended to be the strongest. While rigid XY-parts experienced lowest modulus of elasticity and highest elongations at break, the flexible XY-parts experienced an opposite trend. The investigation into orientation provides the first known research into mechanical behavior of photopolymers containing different concentrationsof TB+ material. Regarding aging effects, ultimate tensile stress of VW+ was shown to increase with time, and elongation at break clearly decreased.

Regarding aging effects, ultimate tensile stress of VW+ was shown to increase with time, and elongation at break clearly decreased. The material properties of VW+ were not altered by lighting conditions. These findings are the first aging effects studied for the material jetting process.The findings from this paper present multiple opportunities for future research. To more clearly assess material performance with future studies, tolerances of the parts should be monitored. Most likely to compensate for the elastomer-like material, the TB+ parts were consistently printed smaller than all of the other parts. Analyzing the extent of this build trend was not in the scope of this paper. During tensile testing, it was common to see parts fail at the neck region. Moore and Williams, who studied fatigue life of VW+ and TB+, ran into this same issue with the TB+ parts [20].

During tensile testing, it was common to see parts fail at the neck region. Moore and Williams, who studied fatigue life of VW+ and TB+, ran into this same issue with the TB+ parts [20]. Tensile specimens are specifically designed with the smallest cross-sectional area along the gauge length to promote failure in this region; however, because the AM process approximates curves with linear segments, the neck region has frequently become the area for failure. While the effects of this have not been thoroughly examined, a tensile specimen design that performs more consistently with elastomer-like materials should be considered. Future work will also aim to expand the study of aging effects in the material jetting process to multiple materials. As end-use parts vary in their material requirements, an understanding of the aging of TB+ and gradient materials that were studied with orientation in this paper could prove beneficial when making material decisions.

Finally, quantifying the relationship between UV-exposure and material properties will be explored in upcoming studies. The scattering nature of light poses the risk for overcuring a part when adjacent layers are scanned. A look into how UV-exposure affects mechanical properties could help evaluate the impact of additional UV-light on part quality. Overall, further investigations into the mechanical properties of AM photopolymers are necessary to continue to improve understanding of the material behaviors so that informed material selections can be made.5. REFERENCES [1] Standard Terminology for Additive Manufacturing Technologies, ASTM Standard F2792- 12a, 2012.[2] Stratasys Ltd. (2015). PolyJet Technology. [Online]. Available:http://www.stratasys.com/3d-printers/technologies/polyjet-technology.[3] J. Kotlinski, “Mechanical properties of commercial rapid prototyping materials,” RapidPrototyp. J., vol. 20, no. 6, pp. 499–510, 2014.[4] K. Puebla, K. Arcaute, R.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceEXPLORING VARIABILITY IN SHAPE MEMORY PROPERTIES OF STEREOLITHOGRAPHY PRINTED PARTS Y. Y. C. Choong*, M. Saeed†, H. Eng†, P. -C. Su*, and J. Wei†Singapore Centre for 3D Printing, School of Mechanical & Aerospace Engineering, Nanyang Technological University, 50 Nanyang Avenue, Singapore 639798 † Singapore Institute of Manufacturing Technology, 73 Nanyang Drive, Singapore 637662AbstractShape memory polymers (SMPs) are smart materials that can change shape and revert to their permanent shape upon external stimulus. Most fabrications of SMPs are based on conventional methods which limit design freedom, hence additive manufacturing presents an alternative to expanding more possibilities for applications.

Most fabrications of SMPs are based on conventional methods which limit design freedom, hence additive manufacturing presents an alternative to expanding more possibilities for applications. In this study, curing process parameters were optimized for printing of photopolymerized thermoset SMPs by the stereolithography process. Tert-butyl acrylate (tBA) and di (ethylene glycol) diacrylate (DEGDA) were copolymerized with variations in crosslinkers to create networks with well-separated transition temperatures (Tg) that varied in a range from 43.6 to 74.1°C. A fold-deployable shape memory test was performed and revealed that free-strain recovery and retention deteriorate with increasing Tg. Nevertheless, the SMPs can undergo at least 20 repeated fold-deploy cycles before failure.

Nevertheless, the SMPs can undergo at least 20 repeated fold-deploy cycles before failure. These results are intended to provide better understandings in processing SMPs via stereolithography, while exploring variability in Tg widens the range of possible applications.IntroductionAdditive Manufacturing (AM) or 3D Printing has attracted significant interest lately due to its great freedom in design in terms of the geometric complexity of the fabricated parts [1]. The development of AM technology over the years has been reviewed intensively, and research interests in this field have been mainly focused on discussing the constraints of AM processes. As this technology advances, it comes to realize that the parts printed are generally static and non- adaptive. But this unique capability can be complementary to shape manipulation via the shape memory programming.

But this unique capability can be complementary to shape manipulation via the shape memory programming. Thus, combining smart materials such as shape memory polymers (SMPs) with 3D printing offers great potential in two aspects: producing SMP devices with practically relevant complex geometries that cannot be easily made by traditional processing methods; more shape variants can be realized for a 3D printed SMP part via shape memory programming. The time-dependent SME offers an additional dimension (i.e., time), leading to the so-called fourth- dimensional printing [2]. 4D printing is now primarily based on polymer-based AM processes where these printed SMPs offer greater flexibility with more degree-of-freedoms and able to withstand significantly larger recoverable strains for shape transformation as compared to metals or alloys [3].

The combination of functionalities with great flexibility in terms of geometric complexity makes the fabricated SMPs more versatile and effective as an active material.In principle, 4D printing can be realized in two ways according to whether they can be printed as a single material or a combination of multi-materials [4]. One of the most common single materials with shape memory properties used in 3D printing is polylactide (PLA) which1807serves as the most popular filament among other materials used in Fused Deposition Modelling (FDM) [5]. PLA can be recognized as a “4D ready” material due to its thermoplastic nature which displays empirical indication of shape memory functionality such that above its melting temperature, the polymer softens upon heating to enable molding and reshaping, but solidifies back once it is cooled. However, FDM produces parts with lower resolution and poorer surface finish as compared to other 3D printing systems.

However, FDM produces parts with lower resolution and poorer surface finish as compared to other 3D printing systems. Their thermoplastic SMP materials also experience creep after prolonged mechanical stress, and hence significantly affecting the long-term performance of the SMPs. On the other hand, 4D printing is widely demonstrated with the use of multi-materials printing since the launch of Stratasys’ Connex multi-material 3D inkjet printing technology. Different material with different properties can be specifically inkjet in different sections of a designed geometry and by utilizing the water-absorbing or thermal-sensitive properties of the materials, the self-assembly process is activated [2]. The thermoset resins from the inkjet printers do not exhibit shape memory properties on its own, but requires a matrix of rigid plastic and elastomeric material in order to demonstrate shape memory properties.

Despite the numerous advantages brought about from multi-material printers to enable fabrication of responsive parts, these multi-material printed parts are prone to material degradation after repeated cycles of folding and unfolding. They are more vulnerable to interface or boundary cracks due to dissimilar materials which can give non-uniform properties such as coefficient of thermal expansion. Hence, this results in shorter fatigue life of the parts and reducing the durability of the parts.Therefore, to overcome the various issues with current 4D printing, our approach is to intrinsic shape memory properties using develop a single stereolithography (SLA) process. In this process, layers of photocurable resin are selectively cured on a platform using a UV laser, layer by layer, to form the 3D object.

In this process, layers of photocurable resin are selectively cured on a platform using a UV laser, layer by layer, to form the 3D object. Material development and fabrication of SMPs using SLA will be more straightforward since SLA has less restriction in material options due to its open build environment and easily accessible resin vat. Inkjet printers are mostly closed systems where they only allow usage of their own proprietary materials, hence materials and parameters cannot be easily changed.

Inkjet printers are mostly closed systems where they only allow usage of their own proprietary materials, hence materials and parameters cannot be easily changed. Moreover, SLA has been known to produce excellent surface finish which can minimize defect-induced failures, hence enhancing the durability of any printed SMPs during repeated thermomechanical cycles.thermoset resinthat hasMaterials and MethodsPreparation of SMP resins: Commercial tert-butyl acrylate (tBA) monomer were synthesized with di(ethylene glycol) diacrylate (DEGDA) crosslinker (10, 20, 30, 40 and 50wt%) and UV photoinitiator Phenylbis (2,4,6-trimethylbenzoyl) phosphine oxide (BAPO) (2wt%). The chemicals were all ordered from Sigma Aldrich and used as received conditions without further purification.

The chemicals were all ordered from Sigma Aldrich and used as received conditions without further purification. The DEGDA crosslinker was first added dropwise to the tBA monomer, subsequently with the addition of photoinitiators in continuous mixing of the solution using magnetic stirring, followed by planetary vacuum mixer at 1900 rpm until the photoinitiators completely dissolved. The syntheses of the chemicals were performed in an ultra-violet (UV) proof environment to minimize pre-photopolymerization.Fabrication of SMPs using SLA: A modified bottom-up scanning SLA (DigitalWax System 029X, Italy) as shown in Figure 1 was utilized and its build platform is immersed into the resin on a transparent base and the resin is cured by laser scanning from below. The bottom-up1808configuration also uses fewer amounts of resins, which makes it more economically efficient in developing SMP resins for SLA.

The bottom-up1808configuration also uses fewer amounts of resins, which makes it more economically efficient in developing SMP resins for SLA. Photoinitiation was induced by a UV solid state laser at a fixed wavelength of 405 nm with computer controlled scanning speed. After scanning each layer whereby the polymer hardens onto the platform, the z-positioning elevator rises to detach from the bottom surface and allow liquid resin to flow in and the process repeats until the 3D object is completely built. The temperature of the printing environment was kept at below the transition temperature of the SMP to prevent the printed SMP from being too soft and gel-like since the SMP is thermally sensitive.Figure 1.

Schematic of bottom-up SLA.Dynamic Mechanical Analysis: Given that the shape memory effect of thermoset SMPs revolves around a temperature range centered at the glass transition temperature (Tg) in which the material is rigid below the Tg and become rubbery when above it, hence the Tg and viscoelasticity of the SMPs were determined using Dynamic Mechanical Analysis (TA Instruments DMA Q800, USA). Samples printed in the shape of rectangular bars with dimensions of 17.5 mm x 11.9 mm x 1.20 mm were placed onto the DMA single cantilever clamping fixture under a dynamic load of 1 Hz with the amplitude set at 25 µm. The samples were heated from 20°C at a heating rate of 3°C min-1. Glass transition temperature which is defined by the tanδ peak and storage modulus in both the glassy and rubbery state can be analyzed from the DMA.Fold Deployable Test: The shape memory properties of the SLA-printed SMPs were measured based on fold-deploy experiments according to [6].

The specimen (100 mm x 10 mm x 2 mm) was heated to deformation temperature, Td (10˚C above Tg) on a heating plate for 5 minutes, then bent into a ‘U’ shape using a mold with an external diameter of 20mm. The maximum bending angle is measured as maxθ which is equal to 180˚. The specimen is cooled down before unloading the external force. A slight spring back might be observed, where the bending angle is recorded as fiixedθ as shown in Figure 2. Lastly, the specimen was reheated for shape recovery and the final endθ . The shape fixity ratio and shape recovery ratio can be bending angle was recorded as calculated as follows:1809Figure 2.

Lastly, the specimen was reheated for shape recovery and the final endθ . The shape fixity ratio and shape recovery ratio can be bending angle was recorded as calculated as follows:1809Figure 2. Schematic illustration of the setup for fold-deployable shape memory test.Shape fixity ratio:Shape recovery ratio:fRrR(%)(%)θfixed θ max θθ − max θ==endmaxResults and DiscussionSynthesis of the SMP resin: The SMP resin was prepared based on thermally induced dual component mechanism, where the thermally induced shape memory behavior is contingent on two components – a soft segment that is responsible for fixing the temporary shape when it is above its glass transition temperature, and a hard segment that is responsible for retaining the original permanent shape.

The tBA monomer acts as the soft segment due to its short chain length which allows higher mobility to allow easy deformation to its temporary shape, while the DEGDA forms crosslinkages that remain thermally stable and locked in position even at high temperatures to retain the permanent shape.Dynamic Mechanical Analysis (DMA): The DMA curves of the SLA-printed SMPs are shown in Figure 3. From Figure 3Error! Reference source not found., the glass transition temperatures (T gs) were defined at the peak of the tan delta curve. It can be observed that all samples possess distinctive T g ranging from 53.9˚C to 74.1˚C and the Tg values gradually increase as the concentration of the DEGDA crosslinkers increases. The addition of higher amount of crosslinkers results in higher crosslinking density within the structure of the SMP chains.

The addition of higher amount of crosslinkers results in higher crosslinking density within the structure of the SMP chains. This leads to a closer structure which hinders the movement of the monomers that acts as the soft segment, hence more energy is required in order to change from its original glassy state to a soft rubbery state for shape transformation.1810[1][2]Figure 3. DMA curves of SLA-printed SMPs with tan δ as a function of temperature.The DMA curves also illustrate the storage moduli of the SLA-printed SMPs being dependent on temperature as shown in Figure 4. When the temperatures are below the T gs of the specimens, the storage moduli remain high, denoting that the SMPs are in the rigid glassy state where the movement of the segments is heavily hindered. The storage modulus of an SMP is also a measure of elasticity of the material, hence with increasing content of crosslinkers, there is an observable decline in the storage moduli of the SMPs.

The storage modulus of an SMP is also a measure of elasticity of the material, hence with increasing content of crosslinkers, there is an observable decline in the storage moduli of the SMPs. Nevertheless, all specimens showed a difference of at least 2 orders of magnitude drop in E’ upon reaching its Tg, which is evident of its shape memory properties since a large and sharp drop in storage modulus in the glass-to-rubber transition zone is necessary for a favorable SMP [7].Figure 4. DMA curves of SLA-printed SMPs with storage modulus as a function of temperature.1811Shape memory properties: The average shape fixity ratios for the SLA-printed SMPs with different concentrations of DEGDA crosslinkers is shown in Table 1. The fixity ratio relates to the ability of the SMPs in fixing its temporary shape upon releasing of the external force and Table 1 displays that the fixity ratio decreases with increase in crosslinker concentration.

Theoretically, the cooling before the release of the external force serves as a kinetic trap to store all strain energy imposed on the permanent shape that changes the conformation of the polymer chains and maintains the system in a high energy state [4]. However, there exists heavy friction among the molecules that resist the deformation and with higher crosslinking density within the SMP structure, the restrictive force becomes much higher which result in loss of some stored energy, hence generating spring back that reduces its fixity capability.Table 1. Measurements of fixity and recovery ratio of SLA-printed SMPs.DEGDA [wt%]10 20 30 40 50fixedθ [˚] 169 167 164 161 154fR [%] 93.9 92.8 91.1 89.4 85.6endθ [˚] 180 157 176 162 162rR [%] 100 87.2 97.8 90 90Figure 5 shows the shape recovery ratio based on the bending angles measured as a function of time.

It can be seen that only the SMPs with 10wt% DEGDA concentration achieves 100% shape recovery as compared to the rest of the SMPs. This can be correlated to the loss of strain energy due to the large spring back with increasing crosslinking density. Since the kinetic trap did not completely store all the energy during the cooling stage, the entropic energy released upon reheating becomes insufficient to fully drive the molecular chains back to their lowest energy state, returning to its original shape. Hence, full shape recovery can only be observed in the printed SMP with lowest crosslinker concentration.Figure 5. Shape recovery of SLA-printed SMPs.1812A series of shape memory cyclic tests was performed on the SMP with 10wt% DEGDA crosslinker to determine the number of folding and unfolding cycles the SMP can withstand before failure by fracture.

A 3D graph in terms of stress, strain and temperature is plotted in Figure 6 to illustrate the cycle steps and the shape memory behavior of the SMP. The SLA-printed SMP on average can withstand up to at least 20 repeated cycles with a maximum of 22 cycles recorded, which is comparable with some commercial thermoset SMPs [8].Figure 6. Shape memory cyclic tests of SLA-printed SMPs with 10wt% DEGDA concentration.ConclusionsIn this paper, single thermoset SMP resins with different concentration of crosslinkers werefabricated using a bottom-up SLA process. Some of the findings have been listed below:1. Distinctive Tgs ranging from 53.9˚C to 74.1˚C are obtained with increasing concentrations of crosslinkers, showing potential for tuneable properties for a wider range of 4D printing materials.2.

Distinctive Tgs ranging from 53.9˚C to 74.1˚C are obtained with increasing concentrations of crosslinkers, showing potential for tuneable properties for a wider range of 4D printing materials.2. Decline in storage moduli of SLA printed SMPs with an increase in crosslinker concentrations due to reduced elasticity, but all specimens exhibit shape memory properties with at least 2 orders of magnitude drop in storage modulus above its Tg.3. Decrease in shape fixity and recovery ratio with higher crosslinker concentrations due to loss in strain energy.AcknowledgementThis project is funded by the Science and Engineering Research Council of SingaporeAgency of Science Technology and Research (A*STAR)-IAP (NTU Grant No. M4070219).References18131.2.3.4.5.6.7.8.ASTM, S., F2792. 2012. Standard Terminology for Additive Manufacturing Technologies. ASTM F2792-10e1, 2012. Tibbits, S., 4D Printing: Multi p. 116-121.

TENSIBLE DIGITAL FABRICATION LANGUAGE FOR DIGITAL FABRICATION PROCESSES Jeffrey I Lipton1, Karl Gluck2, Hod Lipson1,2 Cornell Computational Synthesis Lab 1Department of Mechanical Engineering 2Department of Computer Science Cornell University, Ithaca NY 14850Abstract While additive manufacturing objects are described by the STL and AMF standards, the protocol controlling the fabricator is typically machine-specific. In this paper, we explore a system architecture that converts geometric data into control processes for equipment. We propose a new Extensible Digital Fabrication Language (XDFL) and an interpreted ToolScript language that describes how a geometry is translated into machine commands. An initial implementation of this system architecture was created and deployed as part of the Fab@Home project.

An initial implementation of this system architecture was created and deployed as part of the Fab@Home project. The introduction of a standard process control language will decouple process planning from the equipment manufacturer, thereby catalyzing the introduction of new equipment and development of better process planners.IntroductionAdditive manufacturing has the potential to transform into a horizontal industry. Prior to the personal computer revolution, many companies were vertically integrated, designing all aspects of their technology, from processor to programs. The current state of the computer industry is horizontal, with different companies specializing in parts of the system. Currently in Solid Freeform Fabrication, most devices are vertically integrated, with a single company designing the materials through planning software.

Currently in Solid Freeform Fabrication, most devices are vertically integrated, with a single company designing the materials through planning software. With defined standards for geometric and material interchange formats, SFF could become a horizontally integrated industry.Currently the field of SFF lacks standards for machine commands, and for material and geometric processing. Any standard must address the needs of various communities. SFF is a rapidly growing and changing field with a variety of different techniques and technologies.

Any standard must address the needs of various communities. SFF is a rapidly growing and changing field with a variety of different techniques and technologies. In order for standards to be successful for a wide variety of current and future SFF techniques and technologies, it must address the following concerns:(1) Technology Independence: Given the huge depth and breadth of additive manufacturing techniques, any machine command, and geometric processing standard must easily be able to be adapted to a wide variety of SFF technologies.(2) Simplicity: Any machine command standard must be easy to implement and understand. Commands should be able to be read and debugged in a simple text editor. While this would limit compatibility with low power microprocessors, many SFF systems are computer controlled.(3) Future compatibility: SFF is a rapidly evolving industry, and any command medium and processing standard must be easily extensible.

New features must be easily added aswarranted by advances in technology, while maintaining compatibility with previous versions.Definitions: In this paper, we will use the following definitions of terms:Process – a technique used for SFF such as FDM, Stereo-lithography, Electron Beam Freeform Fabrication (EBF3), etc Machine – a digital fabricator using a specific tool, eg Fab@Home with syringe tool, Fab@Home with valve based tool, Fab@Home with FDM tool, EBF3 machine, Makerbot, RepRap, etcMaterial distribution – a geometry associated with a particular materialBackgroundFor the past three decades manufacturing industries have used G-code for computer numerical controlled processes. G-code was designed for repeated subtractive manufacturing tasks. The language defines tool paths and common machine interactions. Many SFF systems have used this language to contain information about their vectorized paths.

The language defines tool paths and common machine interactions. Many SFF systems have used this language to contain information about their vectorized paths. Industrial and commercial machines such as EBF3 machines and LENS machines use G-code only for pathing information. (1) (2) Low cost kits such as Makerbot and RepRap use G-code to contain information about material deposition, environmental parameters and machine parameters in addition to path information. (3)G-Code has severe limitations on its ability to be a useful command medium in the future. G-code itself has no widely adopted standards or governing body. This has lead to many different companies developing unique standards for their particular machines. In the field of SFF some machines, though using the same technology, have used different G-code language dialects.

In the field of SFF some machines, though using the same technology, have used different G-code language dialects. A Makerbot machine uses M-codes to start and stop an extruder while with RepRap machines the extrusion is treated like an axis of motion in the movement commands. It is difficult to extend G-code. Since each command is distinct and numbered, modifying the function of a command would require generating a new numbered command, or breaking backwards compatibility. G-code is designed explicitly for vectorized interaction and could not easily be used for other types of existing technologies. Additionally it is inherently mono- material, limiting its future usefulness of additive manufacturing.Fab@Home robo-casting systems use a customized XML language called “fab” files inorder to contain vectorized path information. (4) It is an multi-mateiral language, capable of describing a build process of n-number of materials.

(4) It is an multi-mateiral language, capable of describing a build process of n-number of materials. The language itself is flexible but is designed explicitly for vectorized printing using a syringe system. While it is possible to adapt the format for various other deposition heads, it is not a natural process. (5)The SFF industry has used STL as a standard geometry format, and is adopting the new AMF format for geometric and material distribution information. (6) However, for the majority of SFF systems, the processing of material and geometric information is hard coded into print planning programs. This makes specifying unique printing requirements for a given material or geometry difficult. The ability to script the print planning process could provide an easy way to extend the functionality of SFF systems.In finding a solution to the standards model, the web-browser was examined. Two standards are the core of a modern web browser are HTML and JavaScript.

Two standards are the core of a modern web browser are HTML and JavaScript. HTML is human and machine readable, and platform independent. It uses tags to clearly separate data and meta-data. JavaScript is the standard scripting language of web interactions. A variety of engines can securely execute embedded JavaScript code. (7)Specifications: OverviewThe architecture of the proposed system depends on two critical components, the Extensible Digital Fabrication Language (XDFL), and the ToolScript standards. XDFL is an XML-compliant command medium. ToolScript is an extension of the EMCAScript (JavaScript) language. It allows users to script the processing of geometric and material information. ToolScript processes a material distribution, such as an AMF file, along with materials settings, into XDFL.

It allows users to script the processing of geometric and material information. ToolScript processes a material distribution, such as an AMF file, along with materials settings, into XDFL. ToolScript is process-specific, and XDFL is process-specific and materials-specific.AMFGeometricMaterialInformationmappingsToolScriptXDFLAMF FileMaterial SettingsFigure 1: ToolScript processes the materials distribution information from an AMF into XDFL commandsSpecification: ToolScriptThe ToolScript for a given process has three primary functions. Firstly, it must slice and object into slices of given thicknesses. The height of these slices may be constant or vary based on a given geometry. Secondly, it must process the slices into deposition commands. A vectorized process requires the processing of slices into paths. A serial voxelized a process requires the processing of slices into voxels sorted in the order of deposition.

A vectorized process requires the processing of slices into paths. A serial voxelized a process requires the processing of slices into voxels sorted in the order of deposition. A parallel voxelized process requires the processing of slices into bitmaps of regions inside of each slice. Finally, the ToolScript must generate the XDFL and write it to a file. In order to accomplish this, a variety of objects are required by a ToolScript.Any given ToolScript file may only need a subset of the objects described here. A ToolScript needs to have an AMF file representation object, which contains AMF region objects. The AMF regions contain the geometric information for a given material. A slicer object must be able to take AMF regions and slice heights, and convert them into a slice object. A slice object must contain a single outer boundary and zero or more inner boundaries. Each slice has a single material and z height associated with them.

A slice object must contain a single outer boundary and zero or more inner boundaries. Each slice has a single material and z height associated with them. Vectorized processes require a pather object, which can convert a slice and information about the process into paths. Each path is a list ofspecial coordinates with an associated material. Voxelized processes require a voxelizer object similar to a pather, which outputs voxel information. Finally, an XDFL writer object is required which can convert representations of voxels, paths, and slices into XDFL code.Using ToolScript it is possible to write custom implementations of the various objects. For example, if one needs a unique path planner for a specific application, it could be written as part of a ToolScript, provided it interacts with the rest of the tool chain as described. It is also possible to use multiple types of pathers for the same object.

It is also possible to use multiple types of pathers for the same object. This would allow one to make an object with a solid top and bottom, and a hollow core, making it air-tight.

A single ToolScript file may only need a subset. Optional items are in parenthesis.Specification: XDFLXDFL has two top-level tags, which denote the different types of information it contains. The palette tag contains information about the materials used in the printing process. The information contained within the palette tag must be globally accessible when using the XDFL file to execute a print. It contains the abstract information about the different materials to be used in the print process. The commands tag contains a sequential list of commands for the digitalfabricator. The commands listed under the command tag may refer to the global information in the palette tag and may locally overwrite them. A value specified in the commands section is valid for any of its parent tag’s children. Unlike the command medium discussed above, XDFL has a built-in knowledge and description of the print’s volume.

Unlike the command medium discussed above, XDFL has a built-in knowledge and description of the print’s volume. It is designed to be useful for vectorized, voxelized, and stratified processes. A list of all of the XDFL tags and their relationships is in appendix A.XDFL: VectorizedA material in a XDFL file for a vector process has five required tags: path height, width, area constant, speed, and compression volume. These allow the XDFL to be created without having explicit knowledge of how the machine works; rather, it only requires knowledge of how it will deposit material. The flow rate of a path is defined below. For any vectorized process, either the flow rate or the path speed can be fixed for a given resolution (width/height values). The compression volume defines how much material is deposited at the beginning and end of a path. A positive value denotes a process over-depositing initially.

The compression volume defines how much material is deposited at the beginning and end of a path. A positive value denotes a process over-depositing initially. A negative value denotes a process under-depositing initially.Path HeightPath Width(b) Figure 3: The equations governing XDFL path flow rates (a), and a diagram of a given path (b)(a)Paths contain a list of points defining line segments. Each point can contain between two and six coordinates. If paths are contained in a layer tag, then they could have a default value for “x”, “y”, or “z” provided by the layer. Coordinates “u”, “v”, and “w” define rotations around the “x”, “y”, and “z” axis respectively. If a machine has less than the provided number of axes, information contained in the tags is ignored. Appendix B contains an example G-Code file and its XDFL equivalent.XDFL: StratifiedXDFL can be used to define purely stratified SFF processes such as laminated object manufacturing.

Appendix B contains an example G-Code file and its XDFL equivalent.XDFL: StratifiedXDFL can be used to define purely stratified SFF processes such as laminated object manufacturing. In order to describe a stratified print, the materials would contain process specific properties and values, and the commands section would contain exclusively layer tags. These layer tags link to image files of the current layer. The properties of the materials could be used to map between the materials and the image files. The XDFL files and image files could be placed in a zip archive with a unique extension for laminated processes. This would ensure that the file completely described the process. An example XDFL file for stratified processes is in Appendix C.XDFL: VoxelizedXDFL files work slightly differently for serial and parallel voxel machines. A serial voxelized XDFL file would be similar to the stratified and vectorized files, but would contain a sequential series of voxel tags.

A serial voxelized XDFL file would be similar to the stratified and vectorized files, but would contain a sequential series of voxel tags. Each voxel tag would define its location in space. It optionally would contain a geometry attribute, which references, and STL or AMF file for visualizations purposes. A vectorized XDFL file may use the voxel tag to deposit a given volume at a given location. A parallel voxel machine would use path tags to move to a location, and then use the bitmap tag to define where space of voxels will be deposited. An example file is in Appendix D.Implementation and PerformanceIn order to test the performance of ToolScript a processing library and environment were necessary. A digital fabrication application library was created, called libFabApp. LibFabApp contains all of the code needed to run ToolScript and process AMF and STL files. The library processes tool files that have the process specific material settings and ToolScript embedded within.

The library processes tool files that have the process specific material settings and ToolScript embedded within. Using libFabApp, FabStudio version one was created to provide a GUI interface for interacting with the library. The first iteration of the library and studio is designed to work with vectorized print processes. Later versions will allow for voxelized and stratified prints.(a)(b)(c)(d)Figure 4: In order to use the new framework in FabStudio a user must load an object (a), position and scale it (b) assign a material to STL files or unmatched AMF files (c), and send to printer (d)We used FabStudio to generate XDFL files for robo-casting processes. By combining the XDFL with a digital fabricator’s configuration information, it is possible to generate the G-Code needed to operate the machine. A custom python script converted the XDFL file into G-code for a Makerbot and RepRap. The G-code from the script was compared to the XDFLs size in a zipped and unzipped form.

A custom python script converted the XDFL file into G-code for a Makerbot and RepRap. The G-code from the script was compared to the XDFLs size in a zipped and unzipped form. Figure six shows that the XDFL is larger than G-code files when uncompressed, but can be reliably compressed compared to the un-standardized G-Code.Size of Various Command FormatsCompressed RepRapCompressed Makerbotcompressed XDFLRepRap G-CodeMakerbot G-CodeCubeXDFL050100150200Size in KBFigure 6: A comparison of XDFL verses two standard G-Code encodings for FDMFuture Work:of Non-vectorized ToolScript objects need to be added to libFabApp. Implementations the non-vectorized XDFL should be tested on various platforms. There are several useful features which should be built on top of the XDFL-ToolScript framework. The simplest one to implement is the embedding of JavaScript into XDFL files. This would allow material data to be calculated dynamically.

The simplest one to implement is the embedding of JavaScript into XDFL files. This would allow material data to be calculated dynamically. Complex curves could be approximated at machine resolution at runtime. Based on the history of websites embedding JavaScript in HTML this could have a variety of benefits. Following embedded JavaScript in XDFL, a hybrid DOM/SAX model with printing related events would allow a system to perform closed loop SFF.Conclusions:XDFL is a unique command medium, since it is applicable to a wide variety of SFF technologies. ToolScript provides a uniform means of programming geometric processing for SFF technologies. ToolScript and XDFL represent a powerful platform. Its full potential can only be realized if it is refined and adopted across systems and user groups.

ToolScript and XDFL represent a powerful platform. Its full potential can only be realized if it is refined and adopted across systems and user groups. By standardizing the processing of geometries and the command mediums of SFF systems, SFF technology could rapidly develop by allowing effort to specialize and the industry to become more horizontal.Acknowledgements:The authors of this paper would like to thank the members of the Cornel UniversityFab@Home Student project team.Bibliography 1. A Design of Experiments Approach Defining the Relationships Betwen Processing and Microstructure for Ti-6Al-4V. Wallave, T A, et al. Austin Tx : 15th Solid Freeform FAbrication Symposium, 2004.2. Free Form Fabrication ofMetallic Components using Laser Engineered Net Shaping (LENS). Griffith, M L, et al. Austin TX : Proceedings of the Solid Freeform Fabrication Symposium, 1996.3. RepRap Project. Mendel User Manual: RepRap GCodes. reprap.org.

tracting Product Performance by Embedding Sensors in SFF PrototypesRahul Rai, Matthew Campbell, Kristin WoodDepartment of Mechanical Engineering The University of Texas at AustinAbstractSFF has been instrumental in improving the design process by providing designers with prototypes that assist them in the communication of design information and design visualization prior to creating fully functional prototypes. Embedding sensors at key locations within an SFF part to extract further data and monitor parameters at critical locations not accessible to ordinary sensors can help immensely in building functional SFF parts. However, this approach requires data acquisition of information such as temperature and strain values from interiors of products. In this work, the authors propose new techniques for embedding thermal sensors and strain gauges into fully dense DuraForm™ during Selective Laser Sintering (SLS) process.

In this work, the authors propose new techniques for embedding thermal sensors and strain gauges into fully dense DuraForm™ during Selective Laser Sintering (SLS) process. The embedded sensors have been used to measure temperatures and strains. They provide higher sensitivity, good accuracy, and high temperature capacity.1. IntroductionSolid freeform fabrication (SFF) refers to any process which can produce parts directly from a 3-D computer model of the part. SFF processes create artifacts without part-specific tooling or human intervention thus reducing cost. SFF processes use a computer representation of the part based on its geometry. SFF processes are used extensively to generate physical prototypes known as rapid prototypes. These rapid prototypes make excellent visual aids for communicating ideas to customers and can be used for design and testing purposes.Use of physical prototypes in design is being reduced by the use of accurate virtual prototypes.

Virtual prototypes are typically computer models, having no physical presence, which predict the behavior of a full scale, physical part. Physical prototype testing, however, has its advantages. An actual, physical prototype can illustrate the performance of the design in the real world. The physical prototype testing can provide reliable information on the performance and lifetime of a product in service. Virtual models can accurately predict the behavior of the full-scale part if they are continuously updated. Only testing the actual prototype and measuring the state at the desired location can do this. Prototypes with sensors embedded at the location of interest would serve this purpose [1]. In this work, the authors propose new techniques for embedding thermal sensors and strain gauges into fully dense DuraForm™ during Selective Laser Sintering (SLS) process. The embedded sensors have been used to measure temperatures and strains.

The embedded sensors have been used to measure temperatures and strains. They provide higher sensitivity, good accuracy, and high temperature capacity.Instrumented prototypes (SFF parts containing arrays of sensors at points of interest) can lead to fast and accurate updating of virtual prototypes. Deep with structures embedded sensors are capable of monitoring parameters at critical locations not accessible to ordinary sensors. Sensors embedded within a structure add intelligence to the structure, which can be useful in a number of ways. Embedded sensors can simply monitor the structure and make available data for corrective action or they can provide information to redesign the state of the system depending upon the system parameters or external factors. However, this requires real-time acquisition of366information such as temperature and strain values from interiors of products. Embedding sensors at critical points of interest is the ultimate goal of this research.

Embedding sensors at critical points of interest is the ultimate goal of this research. Data collected from physical prototypes with embedded sensors could be used to update virtual models. Successfully instrumented prototypes make virtual prototypes more accurate, allowing engineers to forego building a full scale prototype. Skipping construction of full scale prototypes reduces costs and cycle times associated with bringing a new product to market.Embedding sensors has been carried out in a number of different applications and in a various ways. Work has been done on developing piezoelectric sensors to monitor the state of a planetary gear train in helicopters [2]. The embedded sensors provide an input to a neural network algorithm that includes mechanical strain data, casing temperature and other data [2]. Shape deposition manufacturing (SDM) has been used to manufacture heterogeneous parts, by embedding sensors in a prototype while it is being manufactured [3].

Shape deposition manufacturing (SDM) has been used to manufacture heterogeneous parts, by embedding sensors in a prototype while it is being manufactured [3]. Fiber Bragg Grating (FBG) sensors were embedded in metallic, polymers and ceramic structures using SDM. Embedding context awareness is another such application of instrumented manufacturing of devices. Work has been done on the augmentation of mobile devices with awareness of their environment and situation in context [4]. Other applications like MEMS technology have been applied to realize intelligent turbine engines [5]. In this application an array of passive sensors like pressure sensors, crack sensors, and blade position using strain sensors have been developed.The next section discusses the selection of materials, the sensors and the processes for embedding sensors through SLS process. Section 3 elucidates the experiments related to embedding of thermocouple sensors.

Section 3 elucidates the experiments related to embedding of thermocouple sensors. Section 4 delves on the experiments related to embedment of strain gauge. Final section presents conclusion and future research work.2. Selection criteria 2.1 Selection of MaterialA variety of materials can be used to make prototypes using SLS process. Table 1 lists different materials that are potential candidates for the manufacturing of the prototypes. Other research has shown that the SLS process may be also applied to any material with a viscosity that lowers as heat is applied [6]. This feature will result in the eventual use of ceramic and possibly metallic powders. However, of the seven common materials listed in Table 1, due to availability and cost issues, DuraForm composite seems to be the best choice for manufacturing instrumented prototypes through SLS. DuraForm parts have also been used as functional prototypes.

DuraForm parts have also been used as functional prototypes. In this research project DuraForm has been adopted as the material used to manufacture prototypes.No. No. 1 1 2 2 3 3 4 4 5 5 6 6 7 7Material Material Thermoplastics Thermoplastics Powdered metal Powdered metal CastForm™ PS CastForm™ PS DuraForm™ PA DuraForm™ PA DuraForm™ GF DuraForm™ GF LaserForm™ ST-100 LaserForm™ ST-100 Somos™ 201 Somos™ 201367Table 1. List of materials used in SLS2.2 Different layer access methodsIn order to embed sensor inside a part two approaches could be taken. Either, sensors can be fabricated along with the fabrication of the part or off-the-shelf sensors can be embedded during the prototype manufacturing by SFF process.Single layer accessMulti-layer accessFigure 1. Different embedding methods.This paper discusses the second approach, of which there are two methods single and multi-layer access.

Different embedding methods.This paper discusses the second approach, of which there are two methods single and multi-layer access. Single layer access refers to embedding sensors which are thinner than one layer of the SLS process. This would allow sensors to be embedded during the build cycle. Multi-layer access refers to embedding sensors which are thicker than one layer of the SLS process. Figure 1 shows the difference between the two embedding methods. Multi-layer access requires one of two things. Parts must be built with an appropriate sized cavity for embedding after the build cycle, or systems must be in place to remove excess powder for the sensor to be embedded during the build cycle. Multi-layer access provides more robust sensors but requires significant changes to the SLS process. The smaller sensors of single layer access are fragile and more expensive. However, smaller sensors provide more accurate point data [7].

The smaller sensors of single layer access are fragile and more expensive. However, smaller sensors provide more accurate point data [7]. In this research both layer access methods are considered.2.3 Selection of sensorsThe most common prototyping sensors detect and measure acceleration, displacement, strain, pressure, and temperature. However, the process of embedding these sensors during the construction of a SLS part will demand the adoption of additional constraints for the selection of the sensors. Therefore the selection of sensor should be based upon the SLS process specification. In this section four selection criteria based on SLS process requirement is discussed.2.3.1 Operating Temperature/Survival temperatureThe internal chamber of the Sinterstation is maintained near the sintering temperature for the entire process. The upper bound of the temperature depends on glass transition temperature of the material being processed.

The upper bound of the temperature depends on glass transition temperature of the material being processed. For DuraForm this temperature is around 178o C. Any sensor embedded during the process is subjected to this temperature for possibly hours. It is of the utmost importance that these sensors can survive at temperature up to 200 degree Celsius, otherwise the embedment would be useless. Hence, sensors whose operational temperatures meet the 200 degree Celsius requirement will be preferred.2.3.2 Sensor SizeSLS is a layer based process where layer thickness depends on the material being used. Layer thickness for DuraForm material is 0.004”. Therefore, the primary requirement for single368layer access is sensor size. Single layer access requires that sensors have a thickness small enough to fit within a single layer. Therefore, sensors with too large a cross section are eliminated as not satisfying the definition of single layer access.

Therefore, sensors with too large a cross section are eliminated as not satisfying the definition of single layer access. Sensors within 2-3 thousandths of an inch are still considered because layer thickness can be increased. However, this constraint can be relaxed for multilayer access. Multilayer access can accommodate large sensors. The sensor should also possess long enough lead wires for data accumulation purposes. The dimensions of the lead wires should also comply with the restriction imposed on the sensor size.2.3.3 Sensor positioningThe sensor should be located in a position which will accurately convey the model’s information. In this regard alignment with respect to a given surface and embedment depths are two critical factors. For example in the case of strain gauge, it should be perfectly aligned with the surface on which it is mounted; otherwise the measurements will not be accurate.

For example in the case of strain gauge, it should be perfectly aligned with the surface on which it is mounted; otherwise the measurements will not be accurate. Shallow embedment depths will aid in determining where the sensor is located after sintering; however, deeper measurements will better confirm or contradicts the analytical models (which are used for validation purposes).2.3.4 Sensor Coating and AdhesivesSome sensors will require a coating to survive the physical strain of handling, positioning, and sintering. In other situations, the coating may jeopardize the embedment. The coating may facilitate the securing process by providing a stronger bond between the sensor and the material. In strain gauges, the surface bonding to the substrate material incorporates the use of an adhesive to transfer the information to the strain gauge.

In strain gauges, the surface bonding to the substrate material incorporates the use of an adhesive to transfer the information to the strain gauge. If the part is built with a cavity and then the sensor is mounted then an epoxy may be required to close the cavity.2.3.5 CostIn many applications, if not all, the sensors will need to be disposable. Once embedded in a SLS prototype, the sensors will be properly bonded and secured within the material. Any attempted removal of these sensors will likely damage them in the process. Therefore, a search for inexpensive sensor is an important factor in the selection of sensors.The temperature and strain gauges are compared to determine the sensors which best meet the design criteria. The comparisons are presented in Table 2 and Table 3. Overall, fiber optic sensors compare very well against other sensors. Unfortunately, optical fibers must be eliminated from the selection due to high cost associated with the testing.

Overall, fiber optic sensors compare very well against other sensors. Unfortunately, optical fibers must be eliminated from the selection due to high cost associated with the testing. These costs can exceed conventional thermistors and thermocouples by thousands of dollars because of the expensive data acquisition equipment needed to interpret the output. Both thermocouples and thermistors possess the ideal characteristics for the embedment process. They are capable of withstanding higher temperatures and their size allows embedment within a single layer without the need to create cavities in the part. They are also inexpensive and thus disposable. Thermocouples are slightly better than the thermistors because they are more durable and less dependent on the coatings. However, not all thermocouple will work in the domain. The thermocouple’s material and physical properties must be carefully selected. Specifically, a precision fine wire thermocouple is ideal (Omega™ 8).

The thermocouple’s material and physical properties must be carefully selected. Specifically, a precision fine wire thermocouple is ideal (Omega™ 8). Type E thermocouples (Chromega-Constantan) prove to have high sensitivities in low temperature (0-500 C) [9]. The selected diameter of the lead wire is 50.8 um with a bead diameter of 127um. The lead wires are 30.5 cm long and uncoated. This thermocouple was selected for temperature measurement purposes.369Finding an appropriate strain gauge proves somewhat more difficult. The operating temperatures of most of the strain gauges are below the required 200 C mark. Typical foil strain gauges consist of a thin wire adhered to a substrate material. The inclusion of this substrate material is critical, however, it also adds to the thickness of the sensor. Substrate materials also increase the difficulty of sensor selection because they are not generally required to perform in high temperature environments.

Foil strain gauges were the only strain gauges that satisfy all theCharacterist Characterist ics icsType E Type EBead BeadFlake FlakeDisk DiskRod RodBimetallic BimetallicFluoresce Fluoresce nt ntOperating Operating Temp (In Temp (In degree C) degree C)700 700300 300Not Not applicable applicable with high with high temp tempNot Not applicable applicable with high with high temp tempNot Not applicable applicable with high with high temp temp5 to 300 5 to 30050 to 250 -50 to 250Resolution Resolution (In degree (In degree C) C)0.2 0.20.02 0.020.02 0.020.02 0.020.02 0.021 11 1Size SizeBare bid Bare bid dia 32µm dia 32µmBare bid Bare bid dia 75 to dia 75 to 1000µm 1000µm500µm X 500µm X 500µm 500µmDisk dia 1 Disk dia 1 to 25 mm to 25 mmRod dia Rod dia 500 to 500 to 5000 µm 5000 µmRelatively Relatively large for large for fibre fibre optics optics60-800 60-800 µm µmCost CostLow LowLow LowLow LowLow LowLow LowVery high Very highVery high Very highTable 2.

Strain measuring devices and their characteristics.selection criteria. Foil strain gauges from Vishay Micro-measurements (Model # WK-06- 125AD-350) were selected for the purpose.3. Embedding thermocouple sensorTests are conducted to investigate the feasibility of embedding thermocouple sensors with in a single-layer of an SLS part. K type thermocouples with bead diameters of 0.003 inches are used based on sensor selection criteria. DuraForm™ is heated to roughly 170°C before laser scanning and the laser only raises the temperature of powder a few additional degrees. The powder has a higher absorbtivity than the sensors, thus allowing the sensors can be directly exposed to the laser beam during embedment.Embedding tests are conducted to verify the validity of data collected from embedded sensors and to uncover issues in the embedding process itself. Small cylinders are created in a prototype SLS machine using Duraform™.

Small cylinders are created in a prototype SLS machine using Duraform™. Initially, the surface of a freshly sintered part is presumed tacky enough for the sensor to adhere to the part directly after laser scanning. The surface of the freshly sintered cylinder is found to be solidified. Next, the sensor is placed on the surface of the part after a layer of fresh powder is deposited but before the laser begins scanning. The sensor and plastic powder are both scanned by the laser, making the plastic temporarily molten. The sensor is fused to the part, effectively gluing it in place.A couple of unforeseen problems are observed while placing the thermocouples. Lead wire management presents a large problem. Long wires become tangled in the roller of the SLS machine, requiring short lead wires. Shortened lead wires prevent tangling but new issues arise. Roller movement over the part bends the wires back and forth raising concerns of fatigue.

Shortened lead wires prevent tangling but new issues arise. Roller movement over the part bends the wires back and forth raising concerns of fatigue. Placing the sensor and lead wires flat on the surface of the part also presents a major concern. Two sensors are placed adequately on the part surfaces. A third placement results in one lead wire protruding above the layer with solid plastic, linking it to the part. This results in part shift, an unsuccessful build. Tests are performed on two successfully built cylinders to check the data from the thermocouples.Having completed the first step of the feasibility study, embedding the sensors, data is collected and compared to expected results. Figure 2 shows the setup for obtaining the necessary data. A reading of the temperature is taken at the heating surface by a thermocouple secured to the heating surface. A wand type thermocouple measures the T∞ for the cylinder.

A reading of the temperature is taken at the heating surface by a thermocouple secured to the heating surface. A wand type thermocouple measures the T∞ for the cylinder. The embedded thermocouple then measures the temperature inside the cylinder during steady state conditions. An expected temperature is calculated analytically using a 1D heating model:)( xT−T ( T = − s ∞ where, hP kA cm=) eT ∞−mx[1]371All thermocouples are K-typeT∞Heat sourceTsHeat regulationFigure 2. Set up for heating verificationcA is cross Variable h represents convection coefficient; P is perimeter; k is conduction; and sectional area.

Set up for heating verificationcA is cross Variable h represents convection coefficient; P is perimeter; k is conduction; and sectional area. The values for this model match the physical characteristics of the specimen tested.089 sT 031 T ∞ = 5Wh ==CC2 m K 0.025Wk=mKP cAm 0.040 = 1.267 10 × =4 m−Using these values m =251.3The infinite fin length assumption is valid for mL m = Infinite fin length assumption is usedmL >> 1 251.3(0.075 ) 18.8=The model predicts a temperature of 31°C while 30°C is recorded during the experiment. The error is less than three percent which shows that embedment of thermocouples during the build process was successful. A sample is machined to expose the embedded thermocouple. Figure 3 shows both a photograph and x-ray of the sample. One wire seems to cross the other in the x-ray but the second wire can barely be seen in the photograph. This indicates that the second wire is at a different Z height.

One wire seems to cross the other in the x-ray but the second wire can barely be seen in the photograph. This indicates that the second wire is at a different Z height. Duraform™ is an insulator so a short circuit is avoided. However, this reemphasizes the importance keeping the lead wires separate.372Figure 3. X-Ray and photograph of embedding sample4. Embedding strain gaugeEmbedding strain gauges is an important goal of this research. Foil strain gauges are chosen for their low cost as they are cheaper than fiber optic strain gauges for example. However, embedding strain gauges into SLS parts is more challenging than embedding thermocouples. Strain gauges are physically larger than thermocouples. Accuracy of strain gauges relies on good axial alignment, adding a third degree of freedom to sensor placement; X and Y placement plus rotation in the XY plane. They are also more sensitive to the high temperatures of the SLS process, due in large part to their backing material.

They are also more sensitive to the high temperatures of the SLS process, due in large part to their backing material. These gauges generally adhere to the surface of a structure to measure the strain. An embedded sensor fused to the structure will possibly eliminate the need for adhesive.Initial experiments are conducted with dummy sensors in order to develop a technique for embedding. These dummy sensors consist of brass rectangles matching the dimensions of the gauges to be used. Wires are soldered to the base of the rectangle in order to simulate the lead wires of the sensor. The solder beads are made as small as possible to minimize the thickness of the simulated sensor. Figure 4 is an illustration of the dummy sensors.3730 . 2 2 0 ”0.002”Sketch not to scale0.400”0.013”Solder beadsFigure 4. Sketch of Dummy sensorsTwo experiments are conducted to hone the technique for embedding strain gauges.

2 2 0 ”0.002”Sketch not to scale0.400”0.013”Solder beadsFigure 4. Sketch of Dummy sensorsTwo experiments are conducted to hone the technique for embedding strain gauges. The strain gauges will be used in dog-bone structures so strain can be measured with an extensometer and compared to data from the strain gauge. The dog-bones are built as two separate pieces. The bottom half contains a small cavity so the powder roller does not affect sensor placement. The top half is a solid dog-bone shape. CastForm™ is used due to its low sintering temperature, allowing tests to be run without a cool down and warm up cycle between builds.The first test involves embedding one dummy sensor. The goal of this experiment is to identify any major problems with the embedding technique. The bottom half was constructed using general practices for CastForm™. However, the Sinterstation™ is not allowed to perform the cool down cycle typically performed with CastForm.

The bottom half was constructed using general practices for CastForm™. However, the Sinterstation™ is not allowed to perform the cool down cycle typically performed with CastForm. The machine is stopped and opened after the bottom half is finished but before a new layer of powder can be deposited. Excess powder needs to be removed from the cavity before sensor placement. A vacuum cleaner is used to suck away the excess powder. Unfortunately, the vacuum tends to move or, in some cases, remove the entire part from the part-bed. The part is replaced and the experiment continues. Lead wires from the “sensor” are gently pushed into the powder to prevent shifting caused by “sensor” and/or lead wires. Voids in the powder caused by “sensor” placement and vacuum shifting are filled by powder needed to cover part. The first few layers are not covered by needed powder. However, “sensor” and lead wires did not cause shifting.Inspection of part shows weak bonding but bonding is achieved.

The first few layers are not covered by needed powder. However, “sensor” and lead wires did not cause shifting.Inspection of part shows weak bonding but bonding is achieved. This is important because the use of adhesive is undesirable. The voids caused by the vacuum mishap and the disassembled dog-bone sample are shown in Figure 5. Problems are exhibited but embedding does seem possible based on this test. The dummy sensor shows some adhesion to the part and problems such as shifting and incomplete layer deposition are attributable to the embedding technique.374Figure 5. Vacuum mishap and disassembled sample 1Valuable lessons learned from the initial test are used to implement the second test. Two samples are constructed in order to identify the effect of part/sensor orientation on embedding. One part is oriented parallel to the roller and the other part is perpendicular. Figure 6a shows a photograph of the part orientation.

One part is oriented parallel to the roller and the other part is perpendicular. Figure 6a shows a photograph of the part orientation. Both parts protrude beyond the recommended build circle so warping and/or delaminations are expected. A soft bristled paint brush is used to sweep away excess powder once the bottom is completed. Lead wires are bent downward at the part’s edge and the “sensors” are angled for the roller to encounter the downward edge first. Both methods are employed to prevent part shifting. The “sensors” are placed without incident but voids in the powder caused during placement result in incomplete layer deposition. The build is paused and the powder feed distance is briefly increased.

The “sensors” are placed without incident but voids in the powder caused during placement result in incomplete layer deposition. The build is paused and the powder feed distance is briefly increased. The remainder of the build occurs uneventfully.Voids from “sensor” Voids from “sensor” Voids from “sensor” placement placement placement“Sensors” “Sensors” “Sensors”Part bed Part bed Part bedRecommended Recommended Recommended Recommended build circle build circle build circle build circleParallel Parallel ParallelParallel part Parallel part Parallel partPerpendicular Perpendicular PerpendicularPerpendicular Perpendicular Perpendicular part part partFigure 6. a) Part orientation, b) Build two layoutPost-build inspection reveals the appearance of proper embedding. The parallel part exhibits some shifting which is largely attributed to warping, part shrinkage due to improper heating.

The parallel part exhibits some shifting which is largely attributed to warping, part shrinkage due to improper heating. Both parts also exhibit minor delamination due to extension beyond the recommended build circle, see Figure 6b. Although shifting in the parallel part is largely attributed to warping, the perpendicular part does not exhibit any shifting. Results are promising to the point where real sensors are attempted next. The next experiment uses the actual sensors in two perpendicular parts. The strain gauges are embedded with solder leads on the top and with leads on the bottom in order to evaluate the effect of solder orientation. Since actual strain gauges are used in this third experiment, no adhesives are applied as in the other experiments. The absence of adhesive375allows an automated system to omit the process of applying it, thus reducing and simplifying the operations required from an automated system.Observations are made as the experiment is conducted.

One sensor breaks during handling and prior to embedment. This occurrence illustrates the fragility encountered with smaller sensors. The sensor lead wires prove to be very difficult to shape before embedding. The material and thickness of the lead wires make their stiffness such that the wires are very resistant to shaping for embedding. A lead wire fixture is used to pre-shape the wires. However, the sensors are difficult to fit into the fixture and do not stay in the shape dictated by the fixture. The powder feed distance is increased for the first layers of the second half build. The increased feed distance fills voids created during sensor placement. Sensors are also angled toward the roller so it encounters the downward side of the sensor as it spreads a new layer. The strain gauges do not remain flush to the first layer of powder deposition due to the stiffness of the lead wires.

The strain gauges do not remain flush to the first layer of powder deposition due to the stiffness of the lead wires. Several layers of powder are needed to fully cover the strain gauges since they are not embedded within a single layer. Single layer access proves unmanageable with foil strain gauges. The geometry of the “dummy” sensors matches the actual strain gauges but the lead wires prove to be very different. Foil strain gauge embedding is no longer pursued with this research due the stiffness encountered with the real lead wires.5. Conclusions: The initial experimental results reveal that embedding an off-the-shelf sensor in a prototype manufactured using the Selective Laser Sintering process is possible by both single layer as well as a multi-layer approach. Since, the data from these embedded sensors will be used to construct more robust virtual models, it is necessary that the sensors embedded perform to their optimum level and give accurate results.

Since, the data from these embedded sensors will be used to construct more robust virtual models, it is necessary that the sensors embedded perform to their optimum level and give accurate results. This necessitates that the embedding process is repeatable and accurate, which has not been achieved during the initial experiments. The main reason for non- repeatability and lack of accurate positioning is involvement of human during the embedment process. The repeatability and accuracy could be improved by designing automated sensor handling systems that can position and hold the sensors during the embedment process. References:1. N. Patwardhan, Instrumented Prototyping, Masters Thesis 2002. The University of Texas at Austin.2. E. J. Nidhiry, G. L. Anderson. Diagnosing repairs with embedded sensors, 1997. Army Logistician, 27-29.3. X. Li. Embedded sensor in layered manufacturing, Doctoral Disseration, 2001, Stanford University.4.

traction of Fault Patterns on SLS Part Surfaces Using the Karhunen-Loeve TransformIrem Y. Turner Grad uate Research Asst.Kristin L. Wood Associate ProfessorIlene J. Busch-Vishniac ProfessorMechanical Engineering Department The University of Texas at AustinAbstractTo gain a thorough understanding of the fault mechanisms in SLS machines, we decom pose SLS profile signals into independent features using a novel tool called Karhunen-Loeve (KL) transform. These individual features can then be studied separately to monitor the occurrence of fault patterns on manufactured parts and determine their nature. Analytical signals with known fault patterns, simulating profile measurement signals from SLS parts, are used to determine the suitability of the proposed method. Multi-component patterns are assumed to manifest on SLS part surfaces, resulting from faults in the machine, for example, the roller mechanism.

Multi-component patterns are assumed to manifest on SLS part surfaces, resulting from faults in the machine, for example, the roller mechanism. The results of this work determine the suitability of the KL transform for condition monitoring and extraction of fault-indicating patterns.Fault Patterns on Manufactured PartsDetecting and quantifying faults that occur during manufacturing is necessary to ensure the efficient production of accurate parts. The field of fault detection and diagnosis in manufacturing aims at eliminating the occurrence of faults by continuously monitoring the process, detecting faults, and taking corrective action. In this paper, the focus is on monitoring the condition of surface quality. The surface is measured at regular intervals for the purpose of detecting any degradation on part surface quality.

In this paper, the focus is on monitoring the condition of surface quality. The surface is measured at regular intervals for the purpose of detecting any degradation on part surface quality. Faults or deviations in the dynamics of the manufacturing machine or its submechanisms are assumed to leave a "fingerprint" on the surface of the part being manufactured, which manifest as fault patterns. These are the fault patterns that we seek to detect, quantify, and diagnose, in order to take remedial action if necessary.A fault is defined as the inability of a system to perform in an acceptable manner [9]. Faults typically manifest themselves as deviations in observed behavior from a set of acceptable behaviors. Fault detection is the recognition of an unacceptable behavior; and fault diagnosis is the identifica tion of a component or set of components in the system that causes the fault [9].

Fault detection is the recognition of an unacceptable behavior; and fault diagnosis is the identifica tion of a component or set of components in the system that causes the fault [9]. As part of fault detection, analysts collect data, extract relevant features, and compare these extracted features to a specification of correct or incorrect data [5, 9].The method of feature extraction and selection is a critical factor in detecting the correct fault indicating features from manufacturing signals. Complex signals are best analyzed and processed using signal processing tools. In this research, we aim to develop a unified method to detect and diagnose the correct features in an accurate fashion.The most reliable methods of fault diagnosis utilize signal processing algorithms to extract fault features. Poorly performing fault monitoring and diagnosis systems are common in industry, and they result in frequent false alarms or insensitivity to a legitimate failure condition [10].

Poorly performing fault monitoring and diagnosis systems are common in industry, and they result in frequent false alarms or insensitivity to a legitimate failure condition [10]. It is often difficult to detect the proper features in the presence of random effects and nonstationarities. The575task of diagnosis is often too complex and unreliable because of errors introduced due to the careless selection of signal processing algorithms [10J. As a result 1 in this research, signal processing methods will be investigated to evaluate their suitability as feature extraction methods. Specifically, fault patterns on manufactured part surfaces will be analyzed to determine the character, severity, and origin of faults that result in poor surface quality.Similarly, ongoing research aims at finding the best approach to classify the extracted features, in order to determine the existence and cause of a fault.

This step typically relies on previously known limits and specifications, or previously known fault patterns. The most traditional approach to detect faulty features or patterns is to compare the observed features to previously known features. In this research, we want to develop a diagnosis method that does not rely on previous knowledge. A good example of such a case is a completely new system where there is not enough prior experience and system models for fault classification are not available.

A good example of such a case is a completely new system where there is not enough prior experience and system models for fault classification are not available. This constitutes a very difficult problem, as it requires any type of fault pattern to be recognized, quantified, and diagnosed, without prior knowledge of the expected faults.We propose to develop a fault detection and diagnosis approach composed of signal processing tools that will detect and quantify a fault-indicating pattern (feature), regardless of the type and characteristics of the patterns and the observed signals, and that will diagnose the originating source of the fault, without relying on previous knowledge or known patterns of faults. We aim to do this in the context of monitoring and controlling part surface quality in manufacturing, specifically, in Selective Laser Sintering [12].To satisfy our research goals, we propose a five-step approach, shown schematically in Figure l.

The main two general portions of this approach consist of feature extraction and quantification andMONITORED SIGNAL: Part Surface Irregularitiescg~~~~~:MONITORING OF t-----illIIoi Fault Components 1--.:;llIooi submechanisms ....- -1 FROM SUBMECHANISMS i'E---1 MANUFACfURINGDECOMPOSmON:SIGNAL MEASUREMENT(using KL)(using HOS)(vibrations)MACHlNEManufacturing MachineIMPLEMENTATION OF CORRECITVE MEASURESFigure 1: Feature Extraction and Fault Diagnosis Approach.feature comparison for diagnosis. To obtain a decorrelated decomposition of the measured signal, we need to apply a mathematical transformation to the signal [1, 14]. The Fourier transform is such a tool, resulting in a decomposition into different frequency components based on sines and cosines as basis vectors.

The Fourier transform is such a tool, resulting in a decomposition into different frequency components based on sines and cosines as basis vectors. However, the decomposition is not accurate in the presence of nonstationary features.576An alternative is the Karhunen-Loeve (KL) transform, an orthogonal transform which decomposes the signals into decorrelated components in the form of a few empirical basis functions that contain the majority of the variance in the original data. The most common application of this transform is to detect dominant features in order to reconstruct original signals with lower dimensionality. Recent applications include recognition of faces in vision and turbulent structures in fluids [3, 6, 8]. We want to use this tool in detecting features in manufacturing signals, specifically, surface profiles [IIJ.

We want to use this tool in detecting features in manufacturing signals, specifically, surface profiles [IIJ. Our main purpose is to extract individual patterns and study their characteristics to determine whether they correspond to significant faults, and use these patterns to determine their origin.In particular, we investigate the suitability of the Karhunen-Loeve transform to provide a proper decomposition leading to accurate fault detection, quantification, and monitoring.

The following sections describe the theory and motivation of the KL transform and investigate its suitability based on analytical signals which model multi-component signals and the occurrence of faults.This paper focuses on the first two steps.Karhunen-Loeve TransformOrthogonal transforms, and, in particular, the optimal Karhunen-Loeve transform (KL), other wise known as the Principal Components Analysis in the statistical literature, are used in it variety of signal processing applications [1, 2, 3, 6, 8, 15J. However, they have never been used for the purposes of fault detection and diagnosis, or for the purposes of detecting and quantifying faults on manufactured part surfaces. In this work, we extend the KL transform to the extraction of physical patterns from surface signals; the purpose is to decompose multi-component signals into individual patterns and study their characteristics in order to determine the occurrence, severity, and shape of each fault.

We want to assure that any pattern, including stationary and nonstationary, can be detected and extracted using the KL transform. This information will later be used to diagnose the origin of each individual fault.The KL transform presents a great advantage when investigating the occurrence of faults, especially when the shape of the fault patterns are not known in advance. Orthogonal transforms provide a decomposition of the signal into its high-energy content basis functions, leading to an accurate decomposition picture in the time-domain [1, 7]. Because the KL method decomposes signals into empirically determined basis functions, rather than pre-defined sines and cosines, any fault pattern, including stationary and nonstationary patterns, can be detected in theory using this transform.Theory and MathematicsIn the context of surface profile characterization, the signal being monitored is the surface profile of an SLS part.

A total of M "snapshots" are assumed to be collected at regular intervals, with N points each, to allow for a continual monitoring of the state of operation.We first compute the mean vector Xave = it I:m Xmover the total of M samples collected from the profile; then the deviation or departure y;' = Xm- Xave of each sample signal from the mean is computed. The mean is removed to simplify the mathematical derivation of the covariances [4]. Next, the covariance matrix is computed as follows:M = M L.....t Ym Ym6 1 " -> [> JTm=l577(1)The eigenvectors of C are the basis functions iii, computed from:The eigenvalues Ai are ordered and the relevant features are selected by choosing the first n < iV dominant eigenvalues from the set of solutions, following the feature selection criterion used in the literature (about 90% of the total energy is sufficient for reconstruction [3]).

The coefficients ai are computed by projecting each sample vector deviation y-:n onto the basis vectors iii from:[ .... ]T .... ai == Ym UiFor each deviation Y-:n,there are i == 1. ..n coefficients, where n is the number of principal eigen vectors; there are a total of m 1. ..M deviations (snapshots). From another point of view, for each principal eigenvector ui, i == 1. ..n, there are m == 1. ..M coefficients. The collection of these M coefficients for each eigenvector is called the coefficient vector; hence, there are as many coefficient vectors as there are eigenvectors. Each original sample vector Xm is then reconstructed with lower dimensionality by adding this linear combination to the sample mean:Simulations using Analytical SignalsIn this paper, faults are simulated in the context of parts from a Selective Laser Sintering machine. The roller in the SLS machine is responsible for depositing an even layer of powder on the powder bed [12, 13].

The roller in the SLS machine is responsible for depositing an even layer of powder on the powder bed [12, 13]. The quality ("evenness") of the surface of the top layer of powder will determine, in part, the surface quality of the final manufactured part. In a previous paper, the authors have shown that, with certain types of powder, the roller leaves "chatter marks" on the surfaces of SLS parts [12, 13]. Any undesirable changes in the dynamics of the roller will result in a change in the surface quality of the part being manufactured. It is changes such as these that we want to be· able to detect using the KL transform.In order to test the suitability and limitations of the KL transform, we use analytical signals in which the fault patterns are known. This assures the accuracy of our results, since we know exactly what fault patterns the decomposition should provide, and when each type of fault pattern occurs.

This assures the accuracy of our results, since we know exactly what fault patterns the decomposition should provide, and when each type of fault pattern occurs. We first simulate cases where different faults occur during a manufacturing process, and compare these cases to the normal state of operation. With these simulations, we also want to assure that multicomponent signals are decomposed accurately, whether the components are deterministic, stochastic, stationary, or nonstationary. Deterministic and stationary changes are introduced, as well as nonstationary changes over the entire monitored profile.The "normal" state of operation is assumed to produce a multicomponent signal with 2 si nusoids, in the form of AIsin(FIj) + A2sin(F2j), one high frequency (FI == 0.9 rad/sec) , small amplitude (AI == 1 mm), and the other low frequency (F2 == 0.2 rad/sec) and large amplitude (A 2 == 2 mm), plus random (Gaussian) noise (zero mean and low variance of 0.09).

This situation illustrates a typical operation state, in which rotating components, such as the SLS roller, intro duce a fundamental sinusoidal pattern, plus a harmonic, accompanied with random noise from the machine operation or other conditions, such as the surface texture of the roller. Faults, such as bearing wear or misalignment, introduce either additional harmonics, or a change in the magnitude578(2)(3)(4)Ind. 1 2 3 4 5 6 7 8 9 10Normal Abnormal1a Abnormal 1b Abnormal2a Abnormal 2b Abnormal 3 Abnormal 4 829.55 255.45 685.45 240.15 89.81 74.01 84.81 55.62 52.88 2.68 26.23 2.58 2.81 2.25 2.29 2.14 1.97 2.00 0.00 0.00633.47 530.83 240.76 235.70 2.81 2.39 2.25 2.00 1.73 0.006938.42 5738.31 245.89 235.65 2.86 2.62 2.39 2.06 1.79 0.001006.66 956.76 72.23 57.30 2.59 2.27 2.25 2.00 1.97 0.006232.03 5938.75 70.81 57.61 2.75 2.21 2.08 1.95 1.64 0.001037.00 954.40 618.54 485.89 2.66 2.24 2.12 1.98 1.74 0.00Table 1: Eigenvalues: Normal State vs.

Abnormal Statesof the fundamental frequency component. In addition, offsets or linear trends may be introduced as a result of misalignment in the non-rotating elements. In the following, These abnormal cases, as well as other nonstationary faults, are simulated and compared to the normal state of operation. In the following, the "eigenprofiles" represent the principal basis patterns on the entire surface being monitored. Snapshots are collected at regular intervals to monitor the condition of the surface being manufactured. Significant "eigenprofiles" are derived from a covariance matrix to provide the shape of the dominant patterns on surface profiles; the eigenvalue corresponding to each eigenpro file indicates the effectiveness of each feature to represent the original profile.

The features are the coefficient vectors corresponding to each eigenvalue; these indicate any change in amplitude of a dominant eigenprofile along the snapshots collected in time or in space, corresponding to the weight of each basis eigenprofile in each snapshot. The variance of the coefficient vectors follows the mag nitude of the eigenvalue. There are as many coefficient vectors as there are dominant eigenprofiles. The coefficient vector contains M points, indicating the change over M profile snapshots.Stationary Changes in Deterministic ComponentsIn the first set of simulations, the normal state of operation is assumed to produce a multi component signal with 2 sinusoids and random (Gaussian) noise. Ten snapshots are assumed to be collected over the entire profile; each snapshot has 256 sample points.

Ten snapshots are assumed to be collected over the entire profile; each snapshot has 256 sample points. The following set of simulations test the effect of stationary changes in the deterministic components.Abnormal States 1a and 1bIn the first case, the high frequency component of the signal increases in magnitude, while the second component remains unchanged. Abnormal states 1a and 1b represent two different magni tudes. We are investigating whether changes in the amplitude of one of the sinusoidal components of the multi-component signal can be detected with the KL transform. This might, for example, simulate the case when a sinusoidal pattern on the SLS surfaces, caused by roller chatter [12J, exceeds its normal limits in magnitude, thus reducing the surface precision and accuracy; remedial action will have to be taken to diagnose the cause of this increase.The eigenvalues from abnormal states 1a and 1b, compared to the normal state, are shown in Table 1.

The first observation is that, for the normal state of operation, the multi-component signal is decomposed into 4 significant eigenvalues, summing to 98% of the total energy in the signal. The first two eigenvalues (#1,#2) correspond to low-frequency sinusoidal eigenprofiles, while the next two eigenvalues (#3,#4) correspond to high-frequency sinusoidal eigenprofiles.579For abnormal states 1a and 1b, we also obtain 4 principal eigenvalues and eigenprofiles. As shown in Table 1, when we increase the magnitude of the high-frequency component of our original signal, we notice that the eigenvalues corresponding to the high-frequency component increase in magnitude. The eigenvalues corresponding to the low-frequency component remain approximately the same as in the normal state of operation.These trends are more easily noticeable when we study the shape of the eigenprofiles and the corresponding coefficient vectors.

This result is expected since we are only changing the magnitude of the high-frequency sinusoidal component, and not the frequency. The change in magnitude is detected using the coefficient vectors. Also notice that, the coefficient vectors corresponding to the high-frequency component show a steady increase in magnitude, On the other580250250hand, the coefficient vectors corresponding to the low-frequency component show no detectable change. The change in magnitude is successfully detected by monitoring the coefficient vectors.Abnormal States 2a and 2bIn this case, we simulate an increase in the magnitude of the low-frequency component. The simulation results show that the eigenvalues corresponding to the low-frequency component in crease in magnitude, as the magnitude of the low-frequency sinusoidal component increases.

The simulation results show that the eigenvalues corresponding to the low-frequency component in crease in magnitude, as the magnitude of the low-frequency sinusoidal component increases. The eigenvalues corresponding to the high-frequency components remain approximately the same in magnitude as in the normal state (see Table 1).In addition, the first four eigenprofiles from abnormal states 2a and 2b, are the approximately the same as in the normal state and in abnormal states 1a and 1b. This result is not surprising since we are not changing the shape of the pattern, but only the magnitude. The change in magnitude In this case, the coefficient is be properly detected with the corresponding coefficient vectors. vectors corresponding to the high-frequency component remain unchanged, while the coefficient vectors corresponding to the low-frequency component show a steady increase in magnitude.

vectors corresponding to the high-frequency component remain unchanged, while the coefficient vectors corresponding to the low-frequency component show a steady increase in magnitude. Once again, the coefficient vectors can be monitored to detect the change in magnitude of any frequency component in the measured signal. In addition to these results, further simulation shows that, when both the high-frequency and the low-frequency components of the multi-component signal are increased in magnitude, the changes in magnitude are again successfully detected by monitoring the coefficient vectors (abnormal state 3, Table 1).Abnormal State 4Finally, we investigate whether the KL transform detects the occurrence of an unexpected fault. The fault, in this case, is introduced in the form of an additional frequency component.

The fault, in this case, is introduced in the form of an additional frequency component. This simulates the case where the bearing of the roller in the SLS machine undergoes wear, hence introducing a harmonic, in addition to the fundamental frequency component. The first two sinu soidal components are the same as the normal state of operation; a third frequency component is added (abnormal state 4). In this case, the addition of a third frequency component introduces an additional set of eigenvalues (see Table 1). Therefore, the occurrence of an additional frequency component can be detected by monitoring the number of significant eigenvalues. The accuracy of this case needs to be investigated further.N onstationary ChangesThe second set of simulations tests the effect of introducing sudden and/or gradual nonstation ary changes to the monitored signal.

These cases simulate the occurrence of faults after a certain point in time, introducing a nonstationary pattern over the monitored signal. Four cases are sim ulated, presented next. The eigenvalues for each of these cases, compared to the normal state of operation, are presented in Table 2.Case 1In the first case, signals from the normal state of operation are assumed to have a change in the mean offset after a certain period of time, hence introducing a nonstationary change in the signal that must be detected during monitoring; the first ten snapshots indicate normal status, while the next ten snapshots indicate an offset.

This situation, for example, simulates a sudden change581Index Normal 255.45 240.15 74.01 55.62 2.68 2.58 2.25 2.14 2.00 0.001 2 3 4 5 6 7 8 9 10Case 1 1602.78 254.08 239.45 69.86 55.23 1.66 1.41 1.32 1.24 1.22Case 2 3243.81 3191.27 71.59 56.26 1.52 1.49 1.36 1.33 1.26 1.22Case 3 2253.35 2239.68 1397.22 69.79 56.32 1.20 1.08 1.05 1.01 0.92Case 4 346191.37 254.37 240.01 69.38 56.66 2.36 2.21 2.11 1.92 0.03Table 2: Results: Eigenvalues for Normal State vs. Cases 1 through 4in the vertical position of the SLS part while it is being formed, caused by the residue deposited on the roller mechanism. As shown in Table 2, this fault manifests itself as an added principal eigenprofile, with a relatively large eigenvalue; the remaining eigenvalues are approximately the same as the normal state of operation. The simulation results show that the first eigenprofile resembles a straight line when plotted with the remaining sinusoidal eigenprofiles.

The simulation results show that the first eigenprofile resembles a straight line when plotted with the remaining sinusoidal eigenprofiles. In addition, the coefficient vector corresponding to the first eigenvalue shows the change in the mean offset very clearly, after the tenth snapshot, as shown in Figure 4. As a result, this additional pattern, defined by a straight line, can be detected by monitoring the occurrence of a fifth eigenprofile, while its significance can be evaluated by monitoring the shape of the coefficient vector. Note that no change is observed in the remaining eigenprofiles or coefficient vectors, when compared with the normal state of operation.50 ,.-......,.._-r-_...--..:..co:.;..ell..:..ICIe..:..nl...:,vee:.;,:.lor...;",..;;..:c'T=-se_1- - - ,_ - - , - _ - . - - - - -.

the coefficient vectors corresponding to the low-frequency eigenprofiles show a change in magnitude after the tenth snaphot, and return to their original normal state after the twentieth snapshot. The coefficient vector corresponding to the third eigenprofile indicates the change in the offset value after the twentieth snaphot, while the coefficient vectors corresponding to the last two eigenprofiles show no change.In addition, as expected,Case 4In the final case, the normal state of operation is interrupted by the addition of a linear trend with a positive slope. This simulates, for example, a change in the table or powder bed slope while the SLS part is being manufactured. The results show the addition of a fifth eigenvalue (Table 2, Case 4) of large magnitude, while the remaining four eigenvalues are approximately the same as the normal state of operation.

Changes in the magnitudes of these basis patterns are successfully reflected in the corresponding coefficient vectors. The introduction of additional patterns are also detected as additional basis patterns. For example, in case 1, the KL transform successfully decomposes the signal into its individual patterns, i.e., 2 sinusoids and a straight line. Monitoring is done by detecting a fifth eigenprofile and studying the change in the cor responding coefficient vector. In addition, the nonstationary change in the offset level is reflected in the relevant coefficient vector, hence allowing the effective monitoring of this nonstationary change in the normal state of operation. In case 2, the sudden change in the magnitude of the sinusoidal patterns is successfully reflected in the relevant coefficient vectors, while no additional patterns are extracted.

In case 2, the sudden change in the magnitude of the sinusoidal patterns is successfully reflected in the relevant coefficient vectors, while no additional patterns are extracted. In case 4, the KL transform extracts the additional basis pattern, represented as a straight line, and detects the introduction of a linear trend, which is reflected in the corresponding coefficient vectors as a linear increase in magnitude over the set of snapshots. Further simulations are necessary to assure that nonstationary patterns oVer snapshots (as opposed to patterns over the entire signal) are also detectable with this method. An example is a set of linear trends with differing slopes over each snapshot.Conclusions and Future WorkThis paper presents a study of the suitability of an orthogonal transform, namely the Karhunen LOEwe transform, in decomposing signals measured from SLS part surfaces.

We show that the KL transform can be used to detect stationary as well as nonstationary patterns, providing an accurate decomposition of multi-component signals. In addition, we show that the eigenprofiles and coefficient vectors derived using the KL transform can be used to effectively monitor and detect the occurrence of faults on measured SLS surface signals.Stationary changes, including changes in the magnitudes of one or more sinusoidal components, are reflected as changes in the corresponding coefficient vectors. The introduction of an additional frequency component is reflected by an increase in the number of principal eigenvalues and eigen profiles. Nonstationary changes in the normal state of operation, including a change in the mean offset level, and a change in the slope of the part surface, are reflected by the introduction of ad ditional eigenprofiles and corresponding changes in the coefficient vectors.

A sudden change in the magnitude of a sinusoidal pattern is reflected by an accurate change in the corresponding coeffi cient vector. As a result, the KL transform is shown to be a proper tool in extracting the relevant features and monitoring stationary and nonstationary changes in the normal state of operation, for those cases simulated in this paper. Further mathematical proof of these results will follow.the authors are currently simulating other fault components, such as transients, and other nonstationary fault components, such as a saw tooth shaped pattern. Other transforms such as the wavelet transforms and higher-order spectral transforms are also being investigated as potential candidates to provide an effective and accurate decomposition of multi-component signals.

The results of this work will be used to develop an accurate fault detection and diagnosis method to assure the quality of surfaces from manufacturing machines.In addition to the work presented in this paper,584AcknowledgementsThis material is based on work supported, in part, by The National Science Foundation, Grant No. DDM-9111372; an NSF Presidential Young Investigator Award; by a research grant from TARP; plus research grants from Ford Motor Company, Texas Instruments, and Desktop Manufacturing Inc., and the June and Gene Gillis Endowed Faculty Fellowship in Manufacturing.References[1] A.N. Akansu and R.A. Haddad. 111ultiresolution Signal Decomposition: Transforms, Subbands, Wavelets.Academic Press, Inc., San Diego, Ca, 1992.[2] V. R. Algazi, K. L. Brown, and M. J. Ready. Transform representation of the spectra of acoustic speech IEEEsegments with applications, part I: General approach and application to speech recognition.

Spiewak. A predictive monitoring and diagnosis system for manufacturing. Annals of the CIRP:manufacturing technology, 40(1):401-404, 1991.[11] LY. Turner, R.S. Srinivasan, and K.L. Wood.Investigation of characteristic measures for the anal ysis and synthesis of precision-machined surfaces. Journal of Manufacturing Systems, 14(5):378-392, September/October 1995.[12] LY. Turner, D.C. Thompson, R.H. Crawford, and K.L. Wood. Surface characterization of polycarbonate parts from selective laser sintering. In SFF Symposium Proceedings, pages 181-188. The University of Texas at Austin, August 1995.[13] LY. Turner, D.C. Thompson, K.L. Wood, and R.H. Crawford. Characterization of surface fault pat terns with applicaiton to a layered manufacturing process. Submitted for review to the Journal of Manufacturing Systems, Special Issue in Layered Manufacturing Systems, February 1996.[14] D.J. Whitehouse. Handbook of Surface Metrology.

traterrestrial	
 Construction	
 Using	
 Contour	
 CraftingBehrokh Khoshnevis, Professor (khoshnev@usc.edu) Jing Zhang, PhD Candidate (zhan947@usc.edu) Industrial & Systems Engineering, University of Southern CaliforniaAbstract:Most proposals for construction of settlements on Moon and Mars are based on transporting structural elements from Earth and assembling them at the destination. A far less expensive and potentially practical approach is using Contour Crafting, a large-scale AM process, in conjunction with in-situ materials. Our trials with sulfur based concrete and sintered lunar regolith simulant made by NASA show strong promise. Our project ultimately aims at demonstration of lunar outpost infrastructure construction involving landing pads, blast walls, roads, shade walls and protective hangars.

Our project ultimately aims at demonstration of lunar outpost infrastructure construction involving landing pads, blast walls, roads, shade walls and protective hangars. This paper reports on our very early efforts in the first stage of the project.Keywords: Contour Crafting, Sulfur Concrete, Lunar RegolithIntroduction:As interest in space exploration increases, human beings are to return to Moon and landon Mars in the future. To carry out experiments and live a longer time on extraterrestrial planets, permanent habitats are needed to shield heat, cold, cosmic rays and meteorite impact. Due to negligible lunar atmosphere or scarce Martian atmosphere, spaceship landing and launching off will cause frozen ice crystal from the fuel and gravels flying at a fast speed far away. High speed flying objects including micrometeorites threaten devices and inhabitants around. With solid landing pads and protective walls, landing and launching off will be much safer.

High speed flying objects including micrometeorites threaten devices and inhabitants around. With solid landing pads and protective walls, landing and launching off will be much safer. Furthermore, hangars are needed to protect the landers and other equipment from radiation and micrometeorites and roads are needed to transfer the landers between the landing pads and the hangars.Most proposals for construction on Moon and Mars are based on transporting structuralelements from Earth and assembling them at the destination. Such approach is expensive and infeasible for large-scale implementation. Our approach is based on the use of in-situ material and digital fabrication techniques to construct the infrastructure elements. Intense research has been carried out in in-situ resource utilization (ISRU) and specifically lunar regolith characteristics have been extensively studied.

Intense research has been carried out in in-situ resource utilization (ISRU) and specifically lunar regolith characteristics have been extensively studied. In our research we are exploring the possibilities of using sulfur based concrete and molten regolith as construction materials to be processed by Contour Crafting.250Sulfur concrete is a mixture of pure sulfur with aggregate, which becomes strong as hydraulic concrete when heats up and cools down. Sulfur is extracted from Troilite soil (FeS) on the moon as a byproduct of iron. Toutanji [1] reports experiments, in which 35 % sulfur with 65% JSC-1 lunar regolith simulant by mass is mixed and heated up to 130-140 oC, and then casted into 2-in cubes. The reported compressive strength is 2300 psi at -27 oC, 2500 psi at room temperature. The performance under extreme temperature cycles has also been studied.

The reported compressive strength is 2300 psi at -27 oC, 2500 psi at room temperature. The performance under extreme temperature cycles has also been studied. Unlike hydraulic concrete, no chemical reaction takes place in the hardening process of sulfur concrete, and the material could be recycled indefinitely by heating up to sulfur melting point. Sulfur concrete cures very fast, reaching 90% of its ultimate strength in the first 6 hours by cooling to ambient temperature in terrestrial applications [2].Researchers have also demonstrated building physicalassets by geothermite reaction of lunar regolith and combustive metal powders. Using electrical wires to heat up the mixture and initiate the geothermite, metal powder will react with lunar regolith. Net shape voussoirs for domes have been fabricated in this way [3]. The compressive strength of fabricated parts ranges from 10-18 MPa (1450-2610 psi).

Net shape voussoirs for domes have been fabricated in this way [3]. The compressive strength of fabricated parts ranges from 10-18 MPa (1450-2610 psi). These voussoir pieces could build up a dome structure which is geometrically stable and small cracks in the voussoir pieces will not threaten the whole structure. In this way, about 33 % of aluminum or magnesium powder by mass is needed for construction material.A far less expensive and more feasible way isexpected to be construction by Contour Crafting, a large scale AM process to be used in conjunction with in-situ materials. By using Contour Crafting the need for human intervention in extraterrestrial construction could be minimized, thus reducing the trouble for supplying food, oxygen and safety measures and saving huge amounts by eliminating the need for such resources.

Our trials with sulfur concrete and molten lunar regolith simulant (JSC-1A) show strong promise.Contour	
 Crafting:The basic CC technology works using layering ofFigure 1. Top: Coextrusion by CC nozzle. Bottom: Dome structures built using Contour Craftingpaste material extruded through a nozzle, which is maneuvered by a gantry or a mobile robot. The nozzle is equipped with computer-controlled251trowels and other features that assure controlled surface geometry and possibility of hollow walls with intricate internal structures. The technology also allows integration with other robotic modules for concurrent placement of blocks, tiles and other external features such as plumbing, electrical and sensor modules in and around the extruded structures. The CC technology is currently at Technology Readiness Level (TRL) 6. Successful experimentations have been conducted on several materials, such as plastics, ceramics, composites and concrete.

The CC technology is currently at Technology Readiness Level (TRL) 6. Successful experimentations have been conducted on several materials, such as plastics, ceramics, composites and concrete. The CC technology as a building system tool has targeted the economic viability of the built-up environment and has demonstrated this capability for terrestrial applications. CC machines have produced full-scale structural elements such as walls with complex internal features (Figure 1, top). In fact, NASA has evaluated the CC technology and an elementary CC machine has been used to demonstrate dome construction with internal features at NASA MSFC (Figure 1, bottom) [4].Automated building technology such as Contour Crafting is critical to improving astronaut safety in construction scenarios where conventionally several large and bulky components are proposed employing astronauts in risky EVA procedures.

The assembly and building of roads and platforms are essential to setting up an initial operational capability base.Figure 2. A Contour Crafting robot, housed on an 'ATHLETE' rover, is shown here printing a parabolic vault structure out of processed regolith. The structure is intended to house a lunar lander or other equipment, and is unpressurized. In the background can be seen an array of solar panels intended to supply power to the robot.252Using CC, it is possible to reduce or eliminate astronaut presence on site. We hope to conduct a series of activities to evaluate the merits and limitations of using CC in establishing an Initial Operational Capability (IOC) lunar base. Figure 2 illustrates a CC nozzle assembly and material delivery and melting system mounted on the JPL ATHLETE rover and constructing a lunar hangar by extruding molten regolith.Sulfur	
 concrete	
 application:Sulfur concrete is a relatively new construction material compared with hydraulic concrete.

To use sulfur concrete, aggregates is mixed with sulfur powder and then is heated up to around 130 oC then is cooled down to ambient temperature. The cooled down mix has a compressive strength as high as 17.24 MPa (2500 psi). Sulfur concrete cures faster than hydraulic concrete, and achieves 90% of its final strength within 6 hours. Sulfur concrete is more resistant to acidic and salt and hence it has been made into sewage pipes especially in food processing industries. Sulfur concrete also has better properties under large temperature cycles. The aggregate in our experiments is washed dry sand with a grain size below 1 mm, similar size distribution to that of JSC-1A. Loosely compressed sand has a specific gravity of 1.64g/cm3, and that for regolith simulant is 1.73g/cm3, and for sulfur and sand mixture is 1.68g/cm3.

Loosely compressed sand has a specific gravity of 1.64g/cm3, and that for regolith simulant is 1.73g/cm3, and for sulfur and sand mixture is 1.68g/cm3. Both sand and regolith simulant grains are irregular, and the chemical composition does not affect the binding strength.Sulfur concrete extrusion could realize complete automation, high reliability and flexibility. We have demonstrated that sulfur concrete may be extruded from a nozzle. In the In terrestrial practice heated sulfur concrete mixture is poured into a mold to cool down, which takes time to cool down and move to the construction site to build up the desired structures. The shape of the mold is also limited.

The shape of the mold is also limited. By extrusion, sulfur concrete could be extruded in arbitrary straight or curve forms at the construction site, fulfilling different design objects at a much faster rate and without needing molds.A prototype extrusion system has been built in our lab and the concept is demonstrated.The system contains three main parts, feeding container, flow regulator, and heated extrusion elements. The feeding container stores raw materials and mixes them uniformly at a designed weight ratio. In the prototype system sulfur and aggregate are premixed with a weight ratio of 35% and 65%, respectively. The feeding container is a funnel with the narrow stem around a copper pipe of extrusion usage. While the system is in operation, the mixture flows into the copper pipe by means of vibration.

The feeding container is a funnel with the narrow stem around a copper pipe of extrusion usage. While the system is in operation, the mixture flows into the copper pipe by means of vibration. In the real machine to be developed, a larger hopper will replace the funnel.A flow regulator controls the flow direction of the heated paste mixture and exerts some pressure to the mixture to make it denser and smoother. Currently the flow regulator has a fixed shape and is only suitable for building certain structures of certain shape, like curved walls such as the one shown in Figure 3 is formed. In the new design a flexible regulator compatible to arbitrary shape is expected, and two heated trowels on both sides of the nozzle orifice will be adopted to smooth out the extrudate surfaces.253During the extrusion process the mixture is heated to around 135 oC and is extruded from the copper pipe by the rotation of an auger.

The copper pipe has an inner diameter of 0.96 cm (0.38 inch), and a length of 10.16 cm (4 inch). At the end of the pipe, there is an opening about 0.51 cm (0.2 inch) tall and 1.02 cm (0.4 inch) wide where sulfur concrete could flow through. The opening is connected to the flow regulator. An adjustable heating element of 40 Watts at maximum is closely coupled with the copper pipe and conducts heat to the mixture inside the pipe. An auger with a diameter of 0.95 cm (0.375 inch) rotates in the copper pipe to extrude the mixture out of the copper pipe as shown in Figure 4. A 12 volt gear motor of a ratio of 16:1 drives the auger by a coupler. A vibrator fixed on the auger generates and transmits vibration to the auger as well as to the mixture and the feeding container. The vibration in the auger breaks the arches that can be formed by the bridging effect shown in Figure 3. Without vibration the bridging effect frequently clogs the extrusion barrel.Figure 3.

Without vibration the bridging effect frequently clogs the extrusion barrel.Figure 3. The Bridging effectEach section between two neighboring auger blades (auger pitch) is 1.27 cm (0.5 inch),and contains a mixture with a volume of ! = ! !!! − !!! ℎ = 0.678!"!The heat needed for the sulfur to reach 135C from room temperature 25 oC is!! = 35% ∗ !!"# ∗ ! ∗ 110! ∗ !! + ! = 96.2!The heat needed for the sand to reach 135 C from room temperature 25 C is !!"#$ = 65% ∗ !!"# ∗ ! ∗ 110! ∗ !!"#$ = 21.5!The total energy needed is! = !! + !!"#$ = 117.7!In the equations ro and ri are the outer and inner diameters of the auger, h is the pitch of the auger, !!"# is the specific gravity of the mixture, Cs is the specific heat capacity of sulfur of 0.71J/g/K, and H is the sulfur heat of fusion of 54.0J/g, Csand is the specific heat capacity of sand of about 0.80J/g/K.During the extrusion process, the auger has an angular velocity of about 30 rpm.

The copper pipe has a length of 4 inch, for this section to get out of the pipe, it takes 8 revolutions and 16 seconds. The actual generated heat from the heating element is 640J. The efficiency of the prototype system is 18.4%, which is quite low due to conduction of heat to the environment. The extrusion process is continuous without interruptions. In the lunar vacuum environment, the efficiency could be increased to above 50%, and to process 1 cubic meter of sulfur concrete a solar panel of 7 square meters would be needed to work for only 1 day.Results:The suitability for extrusion of different sulfur concentration is tested. Sulfur concentration of 35% by mass shows feasibility to extrude continuous and high quality walls.

Sulfur concentration of 35% by mass shows feasibility to extrude continuous and high quality walls. Single layer walls of different sulfur concentration have been built as shown in Figure 4.At a sulfur concentration below 30% by mass, friction between the auger and the mixturemight suddenly increase to a degree that the motor could not rotate even under the strongest254vibration. The extruded wall is discontinuous and very loose. At a sulfur concentration above 40% by mass, melted sulfur will flow out from the nozzle, causing the material to be non- uniform and the structure to be weak because of sulfur dominance. For a sulfur concentration of 35%, the extruded process is smooth and the rotation can be maintained at a constant speed.The distancebetween the bottom of the copper pipe and the platform also matters. When the distance is too large the mixture in the copper pipe has no support and will fall down to the platform directly without enough sulfur getting melted.

When the distance is too large the mixture in the copper pipe has no support and will fall down to the platform directly without enough sulfur getting melted. A small distance is preferred if the platform is flat, since concrete flows out in all directions if there is a large gap between the nozzle outlet and the platform.Figure 4. Extruded parts with 40%, 30% and 35% sulfur concentration (left to right)Our measuredcompressive strength of extruded sulfur concrete is 7.79 MPa (1130 psi) at maximum and 3.65 MPa (530 psi) for average. This value is lower than that of the sintered cube in Toutanji’s research. Sulfur concrete in this experiment is poorly pressed. Once a more powerful extrusion system is used to compact the sulfur concrete, the compressive strength is expected to increase significantly.Experiments show that extruded sulfur concrete has promising properties, especially forlunar condition where gravity is 1/6 of that of earth. Here are the main advantages:1.

Here are the main advantages:1. Sulfur concrete gets melted at 120 oC, and the specific heat is low, reducing the energy demand. In practical applications heating the mixture to 130 oC is sufficient for the process.2. Sulfur concrete cures very fast. Experiments show that in about 2 minutes after extrusion sulfur concrete becomes hard and can sustain multiple layers.3. No waste material is generated because all the solid material could be fixed in the sulfur net.4. Material is recyclable. The built part could be recycled by crushing it into small pieces or melted into paste. Material recycling is favored on lunar construction.2555. The interlayer binding is strong.

The built part could be recycled by crushing it into small pieces or melted into paste. Material recycling is favored on lunar construction.2555. The interlayer binding is strong. By building parts in multilayers it is proven that the above layer could adhere to the layer underneath.From the photos of the specimens it can be seen that the fabricated parts contain porosities in the inner structure, mainly caused by expanded air bubbles in the mixture; such phenomena will not exist in the lunar environment where there is no atmosphere.

Compaction immediately after extrusion could also reduce the porosity.Lunar	
 Regolith	
 Sintering	
 and	
 Melting:While sulfur concrete extrusion by CC is suitable for certain lunar structures that will not experience heavy loads and will not be exposed to direct sunlight (which could heat the structure above the melting point of sulfur), processes based on lunar regolith sintering and melting could be used to build landing pad, protective walls and other structures that undergo heavy loads and receive intense radiation by the sun. Sintered plain regolith has high compression strength. To sustain shock impact and tensile stresses, we have discovered that mixing metal powders before sintering significantly enhances the regolith tensile strength. We are also envisioning a variation of the CC method for sintering-based fabrication. Numerous experiments have been carried out to study regolith sintering.

We are also envisioning a variation of the CC method for sintering-based fabrication. Numerous experiments have been carried out to study regolith sintering. The JSC-1A lunar simulant used in our experiments has been obtained from Orbital Technologies Corp. This regolith simulant is designed to be chemically similar to mare regolith. This simulant has a grain size smaller than 1mm, and melts at 1100- 1125 oC.Experimental	
 Setup:Lunar regolith and its mixtures with copper and steel are manually compacted in graphite molds. Compacted regolith mixture is put into a ceramic furnace (CeramPress Qex Porcelain and Pressing Furnace). The sintering temperature ranges from 975 oC to 1100 oC, lasting for 1 hour or 2 hours, in a vacuum condition of 1 torr. Temperature rising rate is 25 oC/min starting from room temperature, and cooling rate to room temperature is decided by heat conduction to the platform in the same vacuum condition.

Temperature rising rate is 25 oC/min starting from room temperature, and cooling rate to room temperature is decided by heat conduction to the platform in the same vacuum condition. The mold is 5.08cm*1.65cm*0.76cm (2 inch*0.65inch*0.3inch) in dimension.Figure 5. Plain regolith sintered at 975 oC for 1 hour256Results:As the temperature increases from 975 oC to 1100 oC, compression strength increases first and then falls down due to porosity near melting temperature. At 975 oC, regolith barely sinters, there is almost no shrinkage, and even a pencil tip could brush off the cooled block, as shown in Figure 5. For plain regolith sintering as shown in Figure 6, from 975 oC to 1025 oC the compression strength of sintered plain regolith increases and then drops down after 1050 oC. The compressive strength at 975 oC for sintered parts is denoted as 0.

The compressive strength at 975 oC for sintered parts is denoted as 0. At 1000 oC, increasing sintering temperature to 2 hours only slightly increases the compressive strength from 6.83 MPa (991 psi) to 10.34 MPa (1500 psi). From 1000 oC to 1025 oC the strength increases rapidly to 36.20 MPa (5250 psi) and 53.5 MPa (7760 psi), respectively. At 1050 oC, sintered block contains large pores, significantly reducing the compression strength. At 1110 oC, regolith melts and the powder swells out of the molds. The final parts contain large voids, which are visible at the surface as shown in Figure 7.Figure 6. Compressive Strength of Sintered Sample with Temperature (The blue and orange dots denote plain regolith)Figure 7. Plain regolith simulant sintered at 1100 oC for 1 hour257Adding steel powder into the regolith could increase the quality of sintered parts at high temperature. The steel powder used is SS316H of 325 mesh size.

The steel powder used is SS316H of 325 mesh size. Without addition of steel powder, plain regolith will swell and voids appear inside the sintered parts. At the temperature of 1050 oC sintered blocks swell out of the molds as seen in Figure 8. With addition of steel powder with a weight ratio of 15%, the mixture has a high compressive stress of 55.57 MPa (8060 psi). The cracked sample after compression test also shows no visible voids in the sintered block.Copper regolith mixture is also tested. Experiments show that copper could also increase the compressive strength but does not help at temperatures as high as 1050 oC. Samples with copper also swell and contain vacancies inside.(a) Plain sintered regolith(b) Sintered regolith steel mixtureFigure 8. Plain regolith and regolith steel mixture sintered at 1050 oC for 1 hourConclusion:CC technology combined with sulfur concrete and regolith sintering has proven feasiblein the lab setup prototype systems.

This approach could significantly save the energy and time used for extraterrestrial construction, and liberates astronauts from the heavy and hazardous work of onsite construction.Experiments have shown the feasibility of the CC technology with sulfur concrete. In-situmaterial of regolith and sulfur could be obtained with ease; saving transportation cost from the earth proves feasibility for space exploration. Sulfur concrete requires much less energy than competitive construction materials and is 100% recyclable. CC will be able to automatically build complex shape structures with sulfur concrete.Compressive strength of sintered plain regolith and mixture could reach 55.16 MPa (8000 psi), which is strong enough for building structures such as landing pads, blast walls and hangers. Sintering of regolith could be possibly used in two ways. First, regolith can be sintered with regular shapes like blocks and voussoirs, and regolith bricks.

Sintering of regolith could be possibly used in two ways. First, regolith can be sintered with regular shapes like blocks and voussoirs, and regolith bricks. Faierson has showed that vousoirs with a compressive strength of 13.79 MPa (2000 psi) could sustain a self weight of 4259 m on the moon, since the dome structure is geometrically stable [3]. In this case regolith brick pavement might be combined with sulfur concrete extrusion. As sulfur concrete is extruded upon the previous layer of bricks, another layer of bricks may be paved upon it with a compression force. Such a combination could greatly speed up the construction rate and stronger buildings could be expected. Second, regolith sintering can be carried out on the construction site to yield the main construction material.

Second, regolith sintering can be carried out on the construction site to yield the main construction material. A variation of the CC process may be used to deliver the regolith mix as a layer onto a selected area and keep it within the confines of three trowels (on sides and258front) and expose the mix to sintering heat for certain time to complete the sintering process while slowly moving the delivery nozzle and trowels forward in a continuous sintering process.Experiments are underway for layerwise fabrication using extrusion of molten regolith. These experiments will be best performed under vacuum as presence of air causes trapped gas expansion under extreme heat that result in voids in the extrudate.Acknowledgement:This research has been performed under a NASA Innovative Advanced Concept (NIAC) Phase I grant.

We have received valuable guidance throughout the project from NIAC program officers and NASA researchers at various NASA centers.References:[1] H Toutanji, B Glenn-Loper, B Schravshuen, Strength and Durability Performance of Waterless Lunar Concrete, 43rd AIAA Aerospace Sciences Meeting and Exhibit; Reno, NV; USA; 10-13 Ja. 2005.[2] Robert E. Loov, Alan H. Vroom. Sulfur Concrete - A New Construction Material, Journal of Prestressed Concrete Institute. 19 (1974) 86.[3] EJ Faierson, KV Logan, BK Stewart, MP Hunt. Demonstration of concept for fabrication of lunar physical assets utilizing lunar regolith simulant and a geothermite reaction, Acta Astronaut. 67 (2010) 38-45.[4] B. Khoshnevis, G. Bekey. Automated Construction Using Contour Crafting--Applications on Earth and Beyond, Proceedings of the 19th International Symposium on Automation and Robotics in Construction. 989 (2002) 489-494.[5] Khoshnevis, B. M. P. Bodiford, K. H. Burks, E. Ethridge, D. Tucker, W. Kim, H.

TRUSION FREEFORM FABRICATION OF BONE-LIKE MINERALIZED HYDROGELS AND MUSCLE-LIKE ACTUATORSPaul Calvert and Zengshe Liu Arizona Materials Laboratories, Tucson AZAbstractExtrusion freeform fabrication has been used to build shapes from agarose, polyacrylamideand polyacrylic acid hydrogels. Contraction and bending can be induced by pH change or application of a voltage between embedded electrodes. Mineral reinforcement can be induced by incorporating salts into the gels and allowing them to react.IntroductionThe growth of organisms normally occurs as the deposition of new material by a layer ofcells at an external or internal surface in an aqueous supporting medium. The layerwise deposition makes the process quite analogous to freeform fabrication. The cells would, however, correspond to a highly parallel multiple head system rather than one with a single moving write head.

The cells would, however, correspond to a highly parallel multiple head system rather than one with a single moving write head. As with some freeforming processes, the actual deposition is a chemical conversion of liquid to solid and the initial deposit may be later transformed to a harder or tougher structure. With this analogy in mind we have been exploring methods by which materials similar to biological tissues might be freeformed.We have thus developed methods to make shapes from water-swollen hydrogels usingextrusion freeform fabrication and have been exploring the application of this method·to the production of synthetic bone and muscle-like actuators.Extrusion Freeform Fabrication of Hydrogels.A hydrogel is a cross-linked solution of a water-soluble polymer with properties similar toJell-o 1. Agarose is a polysaccharide that is used as an electrophoresis medium for biopolymer separation.

Agarose is a polysaccharide that is used as an electrophoresis medium for biopolymer separation. A 5% solution in water melts at about 80°C and solidifies rapidly below 60°C to a strong gel. The solidification process is a phase transition to the formation of helical agglomerates of chains. Using our extrusion system, solids can be written using a heated syringe depositing material onto a cooled. plate.Polyacrylamide gels are much used as a medium for electrophoresis of proteins and arealso used to retain water for houseplants. In this case a solution of acrylamide monomer and about 5% of methylenebisacrylamide crosslinking agent is polymerized by a free radical system. The solution reacts slowly at room temperature but polymerizes in 5-10 minutes at 60°C. Since the monomer solution is very fluid, it is necessary to add a gelling agent to render the solution pseudo plastic and thus allow the parts to hold their shape while polymerization occurs.

Since the monomer solution is very fluid, it is necessary to add a gelling agent to render the solution pseudo plastic and thus allow the parts to hold their shape while polymerization occurs. This is achieved with the addition of 12 wt.% offumed silica as a thickening agent.It is also possible to form gels by writing into a liquid medium. A solution of acrylamide, bisacrylamide, potassium persulfate free radical catalyst and silica is deposited below the surface of an aqueous solution of tetramethylethylenediamine (TEMED) co-catalyst. As the monomer11solution contacts the TE1v1ED solution, the TE1v1ED reacts with the persulfate and initiates the polymerization. To prevent the two liquids from mixing, the TE1v1ED solution is thickened with sodium carboxymethylcellulose. In these circumstances the hydrogel forms in a liquid support medium. It has, however, proved difficult to anchor the hydrogel and stop it drifting during the process.

In these circumstances the hydrogel forms in a liquid support medium. It has, however, proved difficult to anchor the hydrogel and stop it drifting during the process. Osmotic pressure differences tend to lead to swelling of the gel so that good resolution is hard to achieve.Gel musclesMuscle functions by myosin crawling along an actin track driven by the energy ofhydrolysis of ATP which induces a cyclic kinking and straightening ofthe myosin head group. Like a man running on a slippery floor, the force depends on the rate of reaction with ATP and there is no force in the absence of energy input. The electrical stimulation of excised muscle functions by inciting the cells to release ATP, not by directly powering the contraction.Gel muscles depend on the contraction of cross-linked ionic gels as they changetemperature, pH or solvent 2, 3.

These contractions can be large and the forces exerted are comparable to muscle but the response time is very long because water must diffuse from the gel as it contracts. Bending or contraction can also be induced electrically as the field drives ions into or out of the gel. The effect tends to be weak because it depends on establishing an ion concentration gradient within the gel or between the gel and its surroundings. Steep ion gradients cannot be established at the low fields that can be set up in aqueous media. A second problem with these gels is that they generally function immersed in water, which is incompatible with simple machinery.We envisage using freeforming methods to make stacks of gels with interdigitatedelectrodes, both to speed the response and to provide internal reservoirs in which the extruded water can be stored. As a first step toward this, we have formed stacks of six layers of gels using both cross-linked polyacrylic acid and polyacrylamide.

As a first step toward this, we have formed stacks of six layers of gels using both cross-linked polyacrylic acid and polyacrylamide. These have been tested for their response on transfer from acid to base and to field applied between electrodes placed between layers 2 and 3 and between layers 4 and 5.Figure 1 shows the swelling response on being moved from acid solution to base of aseries of 6-layer gels varying from all acrylamide to all acrylic acid via 1+5, 2+4 etc.. Acrylamide gels respond little to pH changes and acrylic acid gels swell strongly in base. It was expected that the mixed gels would curve very strongly as the acrylic acid side expanded. In fact they curved only slightly and the acrylamide portion of the mixed gels swelled equally in the x,y and z directions, driven by the acrylic acid. Apparently the strong swelling of the acrylic acid gel stretches the acrylamide along x and y.

Apparently the strong swelling of the acrylic acid gel stretches the acrylamide along x and y. The concurrent swelling along z is interpreted as a Poisson's ratio of -112 for the acrylamide gel. Such anomalous Poisson's ratios have been previously found for gels 4, 5.Electrical stimulation of similar multilayer stacks is being studied. Figure 2 shows thebending response of six layer stacks with one or two outer acrylamide gel layers either side of two or four acrylic acid gel layers. These samples are run in air with electrodes embedded between the outer acrylic acid layers and the acrylamide. We envision the acrylamide layers as providing a123.

Because the mineral grows in situ in the polymer, it is in the form of fine, highly oriented ribbons, which give bone remarkable stiffness and toughness for a particle reinforced polymer. We hope to freeform similar materials, both for medical applications and as tough synthetic composites.During puberty, the long bones grow by addition of new material to the "growth plate".This is a soft zone extended across the bone near one end. Here a layer of cells continuously forms new cartilage, which has a soft gel-like structure. As the cartilage thickens away from the cells, it becomes mineralized with hydroxyapatite and converts to bone over a distance of 5-10 microns from the cell front.In this way the bone progressively lengthens 6.We hope to produce bone-like composites by a similar process of freeforming gel structures and then inducing mineralization. As a first step, agarose gels were freeformed with a high level of calcium chloride dissolved in the gel.

As a first step, agarose gels were freeformed with a high level of calcium chloride dissolved in the gel. These gels are then immersed in carbonate or phosphate solution to induce mineralization. In initial studies have focussed on carbonates because the precipitation chemistry is much simpler than for phosphate. When the carbonate solution is less concentrated than the calcium chloride in the gel, osmotic pressure drives solution into the gel, which mineralizes, as calcite, with little loss of calcium. In this way up to 14 wt.% mineral can be formed within the gel. While this fraction is too low to give a stiff structure, most ofthe mass of the composite is water and the mineral/polymer ratio is high. On drying, a hard bony structure with up to 70 wt.% mineral is formed, table 1.To induce internal mineralization, multilayer gels have been formed with calcium and carbonate in alternating layers.

On drying, a hard bony structure with up to 70 wt.% mineral is formed, table 1.To induce internal mineralization, multilayer gels have been formed with calcium and carbonate in alternating layers. Precipitation normally occurs uniformly in the layer with the higher initial ionic strength, rather than at the interface. The mineral content is limited by precipitation of the gel structure at high ionic strengths, especially with divalent anions. Weare now reaching higher mineral contents by forming multiple layers containing dispersed powders of slightly soluble minerals, which then reprecipitate as calcium carbonate, leaving a soluble salt.These studies raise a question that is, so far, unanswered. When bone forms, water is lostduring mineralization. There is no obvious physical reason for this because the mineral forms inside the collagen fibers while the water is in a surrounding gel matrix.

There is no obvious physical reason for this because the mineral forms inside the collagen fibers while the water is in a surrounding gel matrix. An understanding of this would help us drive the water out of our gels as they mineralize.ConclusionsA freeformed gel matrix provides a medium within which subsequent or concurrent chemical processes can be used to make new materials with complex microstructures. There are many parallels between this approach and biological growth processes.AcknowledgementsWe would like to thank the Army Research Office for support of this work.14References 1. 2. 3.Dagani, R., "Intelligent Gels", Chem. Eng. News 26-37, June 9, 1997. Osada, Y., Okuzaki, H. and Hori, H., Nature 355, 242-244, 1992. Gong, 1. P., Nitta, T. and Osada, Y. "Electrokinetic modeling of the contraction of polyelectrolyte gels." in Proceedings ofthe Second International Conference on Intelligent Materials. ICIM '94 (1994) (eds. Rogers, C.A.

trusion Freeform Fabrication of Functional Ceramic PrototypesR. Vaidyanathan, J.Lombardi, 1. Walish.Advanced Ceramics Research, Inc., Tucson, AZ 85706S. Kasichainula, Calvert. University of Arizona Arizona material Laboratories, Tucson AZ, 85712Cooper. NASA Marshall Space Flight Center, Huntsville, AL 35812Extrusion Freeforming (EFF) and Fused Deposition Modeling (FDM) processes are established freeforming techniques capable of fabricating complex shaped ceramic prototypes by the sequential deposition and solidification of green ceramic feedstock, layer by layer until the final part results. The freeforming of ceramic parts was accomplished using a commercially available Stratasys 3D Modeler retrofitted with a high-pressure extrusion head designed by Advanced Ceramics Research, Inc. (ACR). The manufactured objects had good dimensional tolerances, as well as real engineering compositions and microstructures.

(ACR). The manufactured objects had good dimensional tolerances, as well as real engineering compositions and microstructures. Ceramic feedstock based on two different silicon nitride powders were developed and successfully used to make prototype parts. Mechanical properties and microstructural characterization of prototype parts were performed.I.IntroductionExtrusion Freeform Fabrication (EFF) is an SFF technique based on the Stratasys FDMTM approacha for the fabrication of functional ceramic prototypes [1]. More details are given in reference 2. While possessing the benefits of Fused Deposition of Ceramics (FDCTM) [1995 Volume from reference 1] and FDMTM, EFF has the added advantages that it can handle higher viscosity feed stock materials and higher extrusion temperatures compared to FDCTM and FDMTM. Similar to other SFF techniques, EFF also allows the sequential deposition of multiple layers to form a complex ceramic shape.

Similar to other SFF techniques, EFF also allows the sequential deposition of multiple layers to form a complex ceramic shape. This has been achieved by retrofitting a high-pressure extruder head to a Stratasys FDMTM modeler (figures 1 and 2). The CAD file is processed by the Quickslice™ software and used to control the high-pressure extrusion head [1]. However, the operation parameters that are automatically set by Quickslice, are optimized for polymer filament type of feedstock. These parameters are not necessarily the optimum parameters for the SFF of ceramic parts. To further control process parameters for SFF of ceramic parts, SML Post™, a Visual Basic based post-processing software was developed to modify selected process parameters. SML Post™ could modify the start-delay, preflow, start-flow, start-distance, main roll-back, flow, and acceleration that were originally set by shutoff-distance, Quickslice™.

SML Post™ could modify the start-delay, preflow, start-flow, start-distance, main roll-back, flow, and acceleration that were originally set by shutoff-distance, Quickslice™. Presently, the road-width, slice-thickness, and fill-patterns are still set by Quickslice. However, SML Post™ could modify these parameters also, if needed.speed,The quality of the EFF green ceramic feedstock has a strong influence upon the robustness of the process and its ability to reproducibly fabricate high strength, dimensionally accurate ceramic components. A. high. degree of homogeneity is desirable in order to minimize density gradients between the binder and ceramic powders.

A. high. degree of homogeneity is desirable in order to minimize density gradients between the binder and ceramic powders. If density gradients are present in the feedstock, it could lead to non-uniform firing shrinkage and formation of defects within thea Stratasys Inc., Eden Prairie, MN327freeform.edceramic bodies3.·The feedstock should also possess a reproducible rheology so that it can be accurately freeformed into thed.esired green ceramic component. Further requirements for the.rheologyofEFF feedstock are a low melt viscosity (extrudable atlow pressures) as well as the.abilitytoundergo fi;l,pidsolidification.upondeposition (enabling mor~ rapid part build rates). The binder should be easily removable from the freeformed green bodies under controlled conditions and leave minimal pyrolysis residue. Finally, the resulting bodies should be readily sinterable into dense ceramic components.Figure 1. Stratasys FDMTM retrofittedFigure 2.

Finally, the resulting bodies should be readily sinterable into dense ceramic components.Figure 1. Stratasys FDMTM retrofittedFigure 2. Close up view of the high pressureWith a high pressure headhead creating a ceramic prototypeIn contrast to EFF, additional requirements must be met for the filamynt feedstock used in FDC processing. In particular, FDC compatible filamentsshould>alsopossesssufficient strength and flexibility such that they can be continuously extruded through a conventional fracturing. Significant FDC filament breakage could lead to Stratasys Modeler without the utilization of dimensional filament feedstock possessing reproducible mechanical properties is an important issue in successfu.l· FDC part fabrication. Therefore, for successful FDC, the filament feedstock needs to m.eet two separate and sometimes divergent sets of property requirements. This is extremely difficult to achieve fornon~idealsystems such as ceramic-loaded binder systems.

This is extremely difficult to achieve fornon~idealsystems such as ceramic-loaded binder systems. The· EFF process avoids this problem by the useof a ceramic feed-rod as the feedstock. This provides the EFF process certain advantages over the current FDC process.inaccuracies and flaws within the FDC parts. Consequently,II. Experimental ProcedureThe entire EFFproCyssinvolves at least five distinct processing stages that result ina structural ceramic part. In addition to the EFF process itself, there are two pre-EFFand two post EFFproces~ing stages.4 The first of these stages is the formulation of a binder system which contains plasticizers, surfactants, and dispersants as well as other additives. More details of the present bindyr system and the rheology of the ceramic-loaded binder system has been discussed in reference 2.

More details of the present bindyr system and the rheology of the ceramic-loaded binder system has been discussed in reference 2. Aftera formulation is decided on, feedrods O.625"x6">length are made by mixing the binder system with the ceramic powder and sinteringaids and then pressing this batch of material into feedrods. The next step is to fabricate a part using the EFF process and burn out the328binder. The subsequent sintering stage densifies the part. Other steps could be added if necessary, such as post-finishing, depending on the quality desired.A.CeramiclBlnder FeedstockDevelopmentThe ceramic feedstock for EFF needs many of the qualities commonly desired in raw materials for ceramic injection molding. 5 Consequently, EFF feedstocks were developed with binder formulations similar to those employed by conventional ceramic forming processes. These formulations consisted of ~ 55 vol% of silicon nitridepowdet dispersed in an organic binder.

These formulations consisted of ~ 55 vol% of silicon nitridepowdet dispersed in an organic binder. The binder is a mixture of polymer, wax, and plasticizer and servesasa vehicle for the freeformed ShN4ceramic powder. The wax is an important component in the binder since it lowers the melt viscosity of the binder polymer at elevated temperatures (ca. > 100°C) while simultaneously enabling the. green body to rapidly solidify and maintaining its dimensional accuracy after freeforming. A liquid plasticizer is believed to be an important binder constituent since it also lowers the binder melt viscosity. Its higher volatility compared to the wax and polymer enables a progressive and more controllable removal of binder components prior to sintering the freeformed c~ramic hodies.6 •. Thesuitability •. o~thisbinder composition· in EFF successfullyextrusionfreeforming and SUbsequent feedstocks was demonstrated.

Thesuitability •. o~thisbinder composition· in EFF successfullyextrusionfreeforming and SUbsequent feedstocks was demonstrated. after pressurelesssintering>97% dense, crack freesiliconnitridepllrts using this type of formulation. Figure 3 depicts both green and pressurelesssintered silicon· nitride· prototypes that Were fabricated using EFF techniques.Figure 3. Phot8~taph of green and sintered siheon nm'me Uli:lI..ll;;Utechniques.Greeniceramic feedstock having.aco11lPosition similar to the values in~able Twas compoundedllsing an Armoloy Prep Mixera •• All raw materials present in the green ceramic feedstock were obtained from commercial sources. The silicon nitride powder was a presSurelessa C. W. Brabender, Hackensack, NJ329sinterable composition containing yttrium and aluminum oxide sintering aids in a 9:3 weight ratio. The saturated elastomer in the binder was an amorphous, noncrystallizable copolymer.

The saturated elastomer in the binder was an amorphous, noncrystallizable copolymer. Gel permeation chromatography of the paraffin wax binder component reported an average molecular weight of 3779 glmole and a polydispersity index of 2.5. The fatty acid plasticizer was a liquid under ambient conditions.Component Silicon Nitride Saturated Elastomer Fatty Acid Ester Plasticizer Paraffin Wax Acryloid AdditiveConcentration (Volume %) z 55 z 25 zlO z5 z5Table I: A typical green ceramic feedstock composition.For rheology characterization, a plunger type Instron rheometer (Model 32ll)c wasutilized for capillary rheometry at temperatures between 120 and 150a C.B. FeedrodProcessing and the EFF ProcessMter shear mixing, the green ceramic material was extruded into feedrods measuring 15.9 mm (5/8") in diameter by 152.4 mm (6") in length. These feedrods were then fed into the barrel of the extrusion head and extruded to produce different 3-D shapes with the EFF process.

These feedrods were then fed into the barrel of the extrusion head and extruded to produce different 3-D shapes with the EFF process. The EFF process itself was optimized by using a Design of Experiment (DOE) procedure? The software-controlled parameters were optimized so that a uniform rate of material deposition was obtained. Figure 4a and 4b show simple parts that were fabricated without optimization of the EFF parameters, while figure 4c shows a simple part that was made after determination of the timization of EFF parameters also improved the on the prototypes. optimized parameters made with different starting silicon nitride powder, such as Figure 5 and 6 are pro Kyocera SN282 and Starck M-11. At present, Alliedsignal AS800 silicon nitride parts are also being made. These results show that the EFF process is capable of handling a variety of starting ceramic powders.a) Figure 4: Cylinders a and b were fabricated prior to optimization.

These results show that the EFF process is capable of handling a variety of starting ceramic powders.a) Figure 4: Cylinders a and b were fabricated prior to optimization. Cylinder c was constructedwith the optimized parameters.b Instron, Canton, MA330Figure 5: Impellers made with Kyocera (left) Figure 6: A turbine blade made withStarck (right) silicon nitride powders.the Starck M-11 powder.Binder Burnout and Sinteringthe organic binder must be removed. The optimum situation is one in which all of the binder has been removed, yet there are no defects created by the removal of the binder system. Thermogravimetric Analysis (TGA) was used to study the compositional loss and degradation of a binder relative to temperature. After TGA is performed, a binder burnout cycle is then devised. Binder burnout cycles are currently being studied so that prototypes with thick cross-sections can be successfully burned out and sintered without fear of producing defects.

Binder burnout cycles are currently being studied so that prototypes with thick cross-sections can be successfully burned out and sintered without fear of producing defects. The current binder burnout and sintering processes can produce parts with greater than 97% theoretical density.Once theprototype has been fabricated usingIII. Results and DiscussionA. CeramiclBinder Feedstock Development and EFF optimizationFormulation modifications were performed in order to obtain lower viscosity and a less shear-rate dependent SbN4 formulation compared to the standard binder formulation. Capillary rheology of EFF feedstock with an acryloid addition for improving the dispersion of the SbN4 powders in the binder system was performed. A comparison of the shear rate-viscosity behavior on standard binder formulations with and without the acryloid addition is shown in Figure 7. The capillary rheology data shows little difference in viscosity at all shear rates.

The capillary rheology data shows little difference in viscosity at all shear rates. However, at the shear rate ranges operative in the EFF process, there was no difference in the viscosity between the EEA binder formulations with and without the acryloid additions. However, the significant finding here was that the extrusion forces during EFF were an order of magnitude lower for the modified formulation compared to the standard formulation. For example, the extrusion pressures for the modified formulation were of the order of 150-200 pounds, while the extrusion pressures with the standard formulation were of the order of 1500 pounds. High extrusion pressures beyond the capacity of the load cell tend to create a situation where the extrusion stops shortly and starts again, creating missed layers and defects in the part.

The high-pressure extruder head is capable of overcoming therhe9logyJilllitations of the EFF formulatiol1s and therefore, other simple indicators ofextrudabilitYIleed to be. developed for the EFFprocess. This could include extrusion force, mechanical properties of flexural bars, as well as ... SEM analysis of fracture surfaces. Room temperature four-point bend tests were performed on flexural bars prepared from EFF formulations with the acryloid addit~qIl' An average strength of 641± 107 MPa was obtained, compared to an average strength of 525 ± 110 MPaJor flexural barsJrom the standard formulation. Figure 8 shows the fracture surface of a test bar. The fra9ture surface did not show any delaminations or missed layers during the freeforming process. figure 9 shows a higher magnification image .ofthe fracture surface from figure 8.Room temperature. flexural tests were. performed on bars produced with the optimized parameters. These results are shown in tableII.

flexural tests were. performed on bars produced with the optimized parameters. These results are shown in tableII. The strength of these bars were comparable with that of previously made bars (polished.~ 613 ± 12, unpolished ~ 594 ± 80)~ Th~se results suggest that optimization oftheEFFparameters provided a good surface finish to the samples.B•.Binder Burnoutand SinteringBinder burnout and sintering currently produce parts that are greater than 97% dense with x-y shrinkages of 18± 3% and z shrinkage of 20 ± 5%. These numbers are similar to numbers obtained from other ceramic freeforming processes8.3321000Figure 8.improvement ofparameters, theprocess hasdemonstrated the capability to fabricate complicated ceramic prototypes using a commercially available Stratasys 3D modeler. Prototypes with excellent surface finish and few defects can be made.

Prototypes with excellent surface finish and few defects can be made. Binder burnout schedules are currently being modified to assure that parts with thick cross-sections can be burned out without warping or defects.T bi II FI a e Raster-build direction 0° 0° 0°/90° alternating 0°/90° alternatingexura strengt Ih db 'leI d' ..fata orf EFF t Surface condition Unpolished Polished Unpolished Polishedf est ars as a unc Ion 0 ras er- UlbftFlexuralstrei594:1:: 80 613:1::12 227:1:: 39 312:1::71f ueClon /"" Jrpa)AcknowledgmentsThis work has been performed under NASA STTR program contract number NAS8-98025.REFERENCES 1. Proceedings of the Solid Freeform Fabrication Symposium, Vol. 1990-1998, Edited by H.L.Marcus, et aI., the University of Texas at Austin, Austin, TX.2. R. Vaidyanathan, J. L. Lombardi, B. Tennison, S. Kasichainula and P.

trusion Freeforming of Nylon 6 Materials John L. Lombardi *, Paul D. Calvert** Advanced Ceramics Research, Inc. Tucson, AZ 85713 * University of Arizona - Arizona Materials Laboratories Tucson, AZ 85712 **INTRODUCTIONNumerous commercial rapid prototyping (RP) processes are capable offabricatingcomplex shaped components. These processes build prototypes "from the ground up" by first reducing a CAD design ofthe desired prototype to a series ofgeometrical slices followed by the precise sequential deposition ofraw material layers upon one another. Unfortunately, these RP processes are limited to producing prototypes from polYmers that exhibit inferior mechanical properties compared to commercial engineering thermoplastics.

Unfortunately, these RP processes are limited to producing prototypes from polYmers that exhibit inferior mechanical properties compared to commercial engineering thermoplastics. Consequently, there are significant advantages in extending the materials processing capabilities ofRP technology into the realm ofproducing tough, high strength functional prototypes from engineering polYmers.BACKGROUNDA possible method offabricating high strength polYmer prototypes utilizes extrusion freeforming techniques (EFF) where parts are built from a series ofprecisely deposited liquid monomer layers that are subsequently thermally polYmerized to form the desired thermoplastic component (1). EFF monomer deposition is typically performed using a sYringe pump fitted with a fine bore nozzle that is interfaced to a CAD based motion control hardware.

EFF monomer deposition is typically performed using a sYringe pump fitted with a fine bore nozzle that is interfaced to a CAD based motion control hardware. A major advantage offreeforming liquid monomer compared to molten engineering thermoplastic feedstock is that the former have lower viscosity (which facilitates accurate dispensing) and do not exhibit many of the viscoelastic rheological problems encountered during extrusion ofthe latter (i.e. die swell) (2). It is necessary that the EFF resin used to fabricate these parts rapidly polymerizes to a high molecular weight thermoplastic in high yield.

die swell) (2). It is necessary that the EFF resin used to fabricate these parts rapidly polymerizes to a high molecular weight thermoplastic in high yield. minimal shrinkage and reaction exotherm otherwise there may be significant stresses or poor adhesion present between the freeformed part layers.It is also necessary that the resin polYmerizes withA review ofthe literature reveals that caprolactam monomer satisfies the above EFF resinrequirements since it is a low viscosity liquid when molten (860 C melting point) and readily undergoes ring opening polYmerization to yield Nylon 6 polYmer. This polYmer is a thermoplastic that is used in a wide variety of structural applications including gears and pump impellers and bearing housings (3,4). In comparison to monomers used in other RP techniques (i.e. stereolithography acrylates) caprolactam also exhibits a lower polYmerization exotherm.

Anionic caprolactam ring opening polymerization reacti mechanismN-Carbamoyl Caprolactam activators are a key component in the aprolactam resin since these compounds lower the activation energy and reaction temperature for Nylon 6 polymerization. Most commercial activators are multifunctional isocyanates which produce Nylon 6 polymer in > 96 % yield within a few minutes after heating the resin above 1500 C (9).EFF Compattble Nylon 6 Resin DevelopmentDespite its numerous advantages, the high reactivity and low shelflife ofcommercialcaprolactam resins precluded their use in EFF processing due to their propensity to gel and clog the EFF apparatus syringe pump· dispensing system. The resin had to be reformulated with an activator that was stable at the EFF pump reservoir temperatures (700 C) while simultaneously having the ability to polymerize the resin shortly after dispensing onto a resin curing platen heated at 1650 C.

Experiments were conducted with various isocyanate compounds to obtain a latent activator. Tetramethylxylxylene diisocyanate (TMXDI) and dimethylbenzyl meta-isopropenyl benzene (TMI) were two compounds found to be suitable EFF compatible latent polymerization activators. The cure properties ofthese two activators are summarized in Table I belQw.474TABLE I TMXDI & TMI NYLON ACTIVATOR CURING CHARACTERISTICSACTIVATORISOCYANATE FUNCTIONALITYTIME (MIN) REQ'D TIME (MIN) REQ'D TO GEL RESIN @ 165°CTO GEL RESIN @ 70°CTMXDI210352045TMI * Nylon resin composition: 30g caprolact~ 1 mol.% activator, 2 mol. % Na Caprolactamate1After determining EFF compatible latent activators, optimal curing conditions for thereformulated resins were detennmed by measuring the percent crystallinity and crystallization undercooling on cast Nylon 6 specimens using differential scanning calorimetry (DSC).

The percent crystallinity within a cast sample varied directly with its modulus and strength while melt undercooling also varied directly with the melt viscosity and molecular weight ofthe Nylon 6 polymer present in each sample (10,11). The results from the DSC characterization of samples cast with TMXDI activator at 165° C for time periods ranging from 15 minutes to 4 hours are presented in Table II below. DSC results for a sample cast with halfthe amount of sodium caprolactamate base and TMXDI activator are also presented in Table II.TABLE II DSC RESULTS FOR Th1XDI CAST NYLON 6Sample (polym Temp-Time)% CrystallinityUndercooling Temp.

T165-154041165-3041.643165-4542.744165-4hL50.444.4165-15 lh TMXDI, Base 47 * Nylon resin composition was similar to that given in Table I above48.5From the above results it can be seen that sample crystallinity and melt undercooling increased with reaction time indicating that the nylon polymer increases in molecular weight and crystallinity upon prolonged reaction. The molecular weight ofthe final nylon polymer was found to level off after a four hour reaction period. Decreasing the amount of sodium caprolactamate base and activator in the resin Yielded a polymer product having both higher molecular weight and crystallinity. This too is expected since the anionic polymerization of caprolactam is believed to be a quasiliving process where polymer product molecular weight is inversely proportional to activator concentration (12). The mechanical properties ofthese samples were also determined using an Instron Model 1011 tensile testing apparatus.

% fume silica, 1 mol. % 1MI, 2 mol. % Na Caprolaetamate32803.5From the results above it can be seen that the extrusion freefonned bars had lower strengths and elongation to break values than those cast under identical conditions. This suggested that the nylon in the extrusion freefonned bars may have a lower molecular weight than that present in the cast bars. It is possible that exposure of nylon resin to the atmosphere during extrusion may have an inhibitory effect upon its polymerization (13).

It is possible that exposure of nylon resin to the atmosphere during extrusion may have an inhibitory effect upon its polymerization (13). The difference in mechanical properties between the extrusion freeformed bar and both the cast control and commercial specimens was not significant, however.477TABLE ill CAST NYLON 6 TENSILE TESTING RESULTSSampleYield Stress (MPa)Young's Modulus (MPa)% Elongation @ Break165-1545.2918563.5165-3043.0518212.4165-4554.0821482.0165-455.623401.5165-15 ~ TMXDI, Base67.4186616.0The results from the tensile testing indicated that both yield strength and sample modulus increased with reaction time. A concomitant decreasing trend in elongation to break values was also observed. This data was supported by the DSC results which showed an increasing degree of crystallinity for samples with heated for long time periods.

This data was supported by the DSC results which showed an increasing degree of crystallinity for samples with heated for long time periods. Furthermore, the increase in sample elongation to break observed in the sample polYmerized with reduced amounts of activator and base suggests that the Nylon present in this sample has a higher molecular weight than that present in the other samples. This too is supported by the DSC results. These results demonstrated that curing the caprolactam resin for four hours at 1650 C produced high strength nylon 6 parts.Extrusion Freeforming ofNylon 6 ResinExtrusion freeforming ofthe resin was accomplished using an AsYmtek Automove Model402 Fluid Dispenser which had a syringe pump attached to the machine dispensing head. In particular, nylon resin was pumped through at 25 gauge needle heated at 700 C onto an aluminum platen heated at 1650 C.

In particular, nylon resin was pumped through at 25 gauge needle heated at 700 C onto an aluminum platen heated at 1650 C. Photographs ofthe AsYmtek Fluid Dispenser and its syringe pump nozzle in the process offreeforming nylon resin are presented in Figure 2 and 3 respectively.Preliminary attempts to accurately freeform the nylon resin onto a platen heated at 1650 Cwere met with some difficulty because the resin was much too fluid at these temperatures and completely wetted the platen surface. Consequently, eight weight percent of hydrophobic, trlmethylsilylated fumed silica thickening agent was added to the resin to increase its thixotropy. The resin was also reformulated with TMI rather than TMXDI activator in order to further increase its pot life. Test bars specimens were then extrusion freeformed using the reformulated resin. Shortly after extrusion freeforming, the test bar specimens were cured for an additional four hours in a mineral oil bath heated at 1650 C.

Shortly after extrusion freeforming, the test bar specimens were cured for an additional four hours in a mineral oil bath heated at 1650 C. After curing, the bars were tensile tested and a comparison was made between these specimens and bars cast under identical conditions. A photograph depicting both a freeformed and a cast Nylon 6 bar specimen is shown below. The tensile testing results are depicted in Table IV below.478..ConclusionsIn this study, it has been demonstrated that a high molecular weight structural thermoplastic can be freeformed from a reactive monomer. Consequently, the materials processing capability ofthe extrusion freeforming technique can be extended into the realm ofprocessing high strength engineering nylon 6 thermoplastic. Furthermore, the freeformed materials possess mechanical properties comparable to nylon 6 produced using conventional processing techniques.

TRUSION METHODS FOR SOLID FREEFORM FABRICATION Paul Calvert!, Robert Crockett!, John Lombardi!&2, John O'Kelly!, Kevin Stuffle2lArizona Materials Labs. 4715 Fort Lowell Rd., Tucson AZ 85712 2Advanced Ceramics Research, 841 E. 47th St., Tucson AZ 85713INTRODUCTIONIt is becoming clear that current SFF methods are members of a potentially very large family. As with current forming methods, it is to be expected that particular materials and objects will lend themselves best to particular methods. Our group has been working for two years on extrusion methods of freeform fabrication where a reactive slurry is extruded into a pattern to form a series of layers that build up a three-dimensional object. A sketch of the apparatus is shown in figure 1.

A sketch of the apparatus is shown in figure 1. The slurry may cure chemically either as each layer forms or in a postcuring oven.As a member of the SFF family, this method offers great versatility in the range of materials that can be formed an in control of the material structure and composition within a part. This paper describes our efforts to extend the application of this method from ceramics to polymer composites, thermoplastics and silica-silica composites.CAD486ProgramComputerGasPressurex-v DriveSlurry in Syringe+PartHotplateFigure 1: Extrusion SFF Equipment50EXTRUSION SYSTEMS ACR SystemsThe ACR extrusion freeforming equipment comprises a computer-controlled, stepper motor driven and screw-actuated gantry plus a vertical linear slide table. A small positive displacement extruder is mounted on the vertical slide. The extruder is a small stiff piston and cylinder assembly.

A small positive displacement extruder is mounted on the vertical slide. The extruder is a small stiff piston and cylinder assembly. The piston is actuated by a second linear slide table driven by a stepper motor. Piston and cylinder were originally glass syringes. These have been replaced with epoxy potted polycarbonate syringes and aluminum pistons. The nozzle (needle) is a shortened Luer Lok hypodermic needle.For polymer extrusion, high extrusion pressures increase greatly the weight of the head. To avoid the problems of moving such a large mass a system has been constructed based on a fixed extrusion head and a moving sample table. A computer-controlled massive vertical extruder is mounted above a three-axis table with two orthogonal cross-tables mounted on a vertical stage beneath the extruder.

A computer-controlled massive vertical extruder is mounted above a three-axis table with two orthogonal cross-tables mounted on a vertical stage beneath the extruder. The outlet nozzle of the extruder is a heated block with a machined internal reduction.AML Dispensing SystemMotion control and slurry placement is provided by an Asymtek model 402 benchtop fluid dispensing system. This unit provides an x-y travel of 12" x 12" and z travel of 3.5" with a step resolution of 0.001" on all axes. The maximum head speed is 12 inches/sec, although dispensing is usually limited to 1-4 inches/sec.Dispensing is through a pinch tube valve connected to a 50 cc syringe. The syringe can be replaced by a feed tube connected to a stationary cartridge allowing a larger supply of slurry.

The syringe can be replaced by a feed tube connected to a stationary cartridge allowing a larger supply of slurry. The slurry is pressurized to 5-20 psi from a nitrogen bottle and is dispensed though an 18-25 gauge needle.Dispensing variables for this system include: Fluid Pressure Tip Gauge Head Velocity Fill Pattern resolution Startup delay Shutoff DistanceThe startup delay (time between the opening of the valve and when the head is instructed to move) and shutoff distance (length before the end of a move at which the valve is instructed to shut off) are required to allow the movement of residual fluid in the needle, to avoid blobbing at line ends.Control is via an IBM PC that handles all aspects of modelling and dispensing in a single integrated package running under Autocad Release 12.

The Autocad AME package provides modelling and object slicing, while software routines written in C provide complete control of dispensing operations, converting Autocad data into Asymtek motion commands and issuing machine control statements.51Drawings of an object are prepared as 3-dimensional solid models in Autocad. To fabricate the object, the user manually sets the dispensing tip to the height of the dispensing surface. The control software reads this value, calculates the correct elevation to slice the model and creates a 2-dimensional section. The perimeter of the resulting pattern is dispensed first, then it is filled in a snaking line pattern.

The perimeter of the resulting pattern is dispensed first, then it is filled in a snaking line pattern. After the layer is dispensed, the user is prompted to reset the needle to the new surface height and the process is repeated.The feedback in the process ensures that the object is fabricated to the highest resolution possible on a layer-by-Iayer basis, without being subject to additive effects of small changes in dispensing thickness. In principle, surface-sensing equipment could provide this information. Current layer thicknesses are 0.02-0.03 inches.One promising filling strategy is to dispense the perimeter of three successive layers before filling the interior. The three layers (0.05-0.1" total height) act as a dam, allowing the fill pattern to be dispensed in a much thicker layer without affecting the overall surface resolution.

The three layers (0.05-0.1" total height) act as a dam, allowing the fill pattern to be dispensed in a much thicker layer without affecting the overall surface resolution. This significantly reduces the total dispensing time and produces flatter, more homogeneous filled regIons.MATERIALS Ceramic SlurriesExtrusion of slurries of alumina in acrylic monomers was described at the same meeting last year. Slurries could be made with solids contents up to 55% and viscosities of 200 poise. The final parts had good fired density and flexural strength [1].Filled ResinsIn the same way, the slurry cured as each layer was extruded. Table 1 shows the properties of a series of silica-filled acrylics. Ceramic green bodies do not need significant mechanical strength and heavily cross linking acrylates, such as TMPTA could be used [1]. In the case of composite materials, the high volume fraction of resin places a higher value on good resin toughness.

In the case of composite materials, the high volume fraction of resin places a higher value on good resin toughness. Consequently the highly cross-linked resins based on TMPTA are unsuitable and were replaced by the more flexible urethane aerylates.Silica-filled acrylic resins could be formed by the same method.Many acrylate and other monomers were investigated before we found a good combination of rapid curing and toughness. Mixtures of a urethane acrylate, hexanedioldiacrylate and N vinylpyrrolidone were found to give rapid curing (less than 1 minute) and good properties. The resin was reinforced with silica particles plus 1.7 vol.% fumed silica which provided enough viscosity that the larger particles did not settle in the slurry.We believe that the toughness arises from the polarity of the urethane which allows good bonding to the filler particles.

It can be seen that the modulus increases by a factor of three as the filler loading increases to 30 vol.% and that the strength remains high. In many filled polymer systems, where there is poor bonding to the filler, the modulus increases but the strength declines with increasing filler content.Fiber-reinforced compositesIn addition to particulate fillers, short fiber fillers can also be used. As shown in Table521, 7% carbon fiber (Hercules AS4), chopped to a length of about 100 microns gives a higher modulus but there is some loss of strength. The strength is almost certainly sensitive to the interfacial bonding. Good fiber-resin bonds can be difficult to achieve with carbon fiber.In situ compositesOther work in our group has been concerned with in situ composites where the reinforcing phase grows within the polymer matrix during or after processing.

For instance, liquid titanium alkoxides have been blended into polypropylene and co-extruded to form a flexible polymer film. Subsequent treatment of this film in a water bath hydrolyzes the alkoxide and precipitates titania to produce a composite with a very fine titania reinforcement.Tetraethoxysilane (TEOS) and water, at a 1:4 molar ratio, have been blended with the acrylate monomers to form an extrudable fluid. This both cures and partly hydrolyzes as the layers are deposited to form a composite. A 24-hour steam treatment of the final part was used to complete the hydrolysis to silica. The parts listed in table I contained 7 vol.% silica from TEaS, pIus 15 vol.% of silica powder and 1.66 vol.% of fumed silica. The properties are similar to the equivalent 22% silica conventional composite.These parts show a volume shrinkage of 15% on reaction. For most of these types of in situ composites, the shrinkage is a serious problem in forming anything but films and fibers.

For most of these types of in situ composites, the shrinkage is a serious problem in forming anything but films and fibers. SFF methods can offer a route to exploitation of this new class of materials since much of the shrinkage can be confined to each layer as it forms, rather than requiring accommodation of the whole part.Table I: Flexural Tests on CompositesCompositionFlexural Strength, MPaFlexural Modulus, GPaResin38.30.96Resin + 3.8 vol% Silica42.81.03Resin + 7.5 vol% Silica48.81.15Resin + 11.2 vol% Silica54.91.39Resin + 15 vol% Silica58.21.56Resin + 22.5 vol% Silica75.12.56Resin + 30 vol% Silica61.52.89Resin + 7 vol% Silica from TEaS50.42.10Resin + 5 vol% Carbon Fiber27.01.2453Ceramic/Ceramic slurriesAs part of a program tied to a future lunar base, where polymers would not be available, we have been making ceramic-bonded ceramic materials.

A slurry of silica powder, ethylsilicate binder and water is formed in ethanol. Ethylsilicate is related to tetraethylorthosilicate (TEOS) but has been partly hydrolysed under acidic conditions to form linear oligomers. Further hydrolysis and condensation lead to silica which will act as a matrix phase in the final sintered body. The ethanol from the slurry and that released on hydrolysis of the ethylsilicate will be recycled by reaction with silicon derived from lunar sources. The water can also be recycled. Essentially this becomes a process for reforming lunar silica into functional parts.The hydrolysis and condensation of ethylsilicate can be acid or base catalyzed. Adding catalysts to the slurry produced uneven results as condensation continued in the syringe before dispensing.Ethylsilicate, water, ethanol and silica are ball milled with alumina balls for 24 hours and then the slurry is poured into the syringe.

Particle loadings of 55-58 vol.% dispense well through a 23-gauge needle at 10-15 psi fluid pressure. After dispensing the layer is briefly treated with a heat gun to remove some ethanol and then is sprayed with aqueous ammonia.The greenbodies are dried slowly and then fired at 1500°C in air. Fired densities have so far only reached 60% of theoretical and fluxes are being explored to enhance the sintering kinetics.Thermoplastic ResinsNylon 6 is a tough engineering thermoplastic that is used for many small parts such as gears, cams and bearing housings. For larger parts the high melt viscosity of Nylon causes problems in filling the mold with polymer of high molecular weight. There has been much development work on reaction injection molding (RIM) methods where molten caprolactam monomer is cast or injection molded and polymerized in the mold. The polymerization can be completed in a few minutes.

The polymerization can be completed in a few minutes. The low viscosity of the monomer allows large molds to be filled and the high molecular weight product has excellent toughness [2,3]. This low viscosity, rapid polymerization system could be readily adapted to SFF extrusion processes.Molten caprolactam, at l20°C, is pumped through a needle onto a heated platen to cause polymerization. A strongly basic catalyst, sodium caprolactamate, and an activator, N acetylcaprolactam, initiate the ring opening polymerization [4]. The reaction occurs rapidly above 150°C and typically produces polymer in 95-98% yield.----'. L-NH-CH -CH -CH -CH -CO~ 2222 J.... LnCaprolactamNylon 654CONCLUSIONSExtrusion methods of SFF allow considerable versatility in terms of the types of materials that can be formed. Apparatus to carry out this technique is simple and readily constructed.

Apparatus to carry out this technique is simple and readily constructed. As well as being a manufacturing method, it provides a useful tool for the materials engineer to test the formation of novel materials. In particular the layerwise formation process permits chemical production of materials to final shape without the normal limitations imposed by slow diffusion rates and large shrinkages.ACKNOWLEDGEMENTSAML would like to thank NASA for support through the UA SERC. ACR would liketo thank DARPA and NASA for support through their SBIR programs.REFERENCES1.Kevin Stuffle, Anthony Mulligan, Paul Calvert and John Lombardi, "Solid freebody forming from polymerizable slurry" Solid freeform fabrication symposium proceedings, Univ. of Texas, Austin, 1993, 60-632.R. Puff, "Lactam based polyamides" Vol. II, CRC Press, Ann Arbor MI, 1991 pp.145-1623.C. Macosko, Plastics Engineering, April 1993, pp. 21-254.1. Sebenda, J. Macromol. Sci. Chem. A6 1145 (1972)55.

TRUSION-BASED ADDITIVE MANUFACTURING OF THE MOISTURE-CURED SILICONE ELASTOMERY. Jin,* †, J. Plott†, A. J. Shih†Department of Mechanical Engineering, Zhejiang University, Hangzhou, China 310027 † Department of Mechanical Engineering, The University of Michigan, Ann Arbor, MI 48105AbstractThe extrusion-based additive manufacturing (AM) of moisture-cured silicone elastomer for complex freeform shape is studied. Due to its low elastic modulus and poor shape retaining ability during the layer-by- layer process, silicone elastomer AM is technically challenging. The experiment for extrusion of room temperature vulcanization silicone elastomer is conducted to study effects of air pressure, nozzle size and speed, layer height and distance between silicone lines on the flow rate and cross-sectional geometry of silicone elastomer AM. The COMSOLTM Multiphysics simulation using the level function to track the silicone-air interface is applied to model the silicone flow.

The COMSOLTM Multiphysics simulation using the level function to track the silicone-air interface is applied to model the silicone flow. Modeling and experimental results of the diameter and flow rate of silicone under the free flowing condition has good agreement and shows the potential for model-based guidelines for AM of silicone elastomers.

Effects of the nozzle speed, layer height, and distance between two adjacent lines are investigated and demonstrate the feasibility and limitations of AM of silicone elastomer.Nomenclaturea b c Distance between two adjacent silicone lines d Diameter of extruded silicone di do Outer diameter (OD) of nozzle tip l Length of the extruded silicone M Mass p Hydrostatic pressure P Air pressure inside the syringe chamber Q Flowrate t v γ ρ Density μ ViscosityLong radius of the cross-section of extruded silicone Short radius of the cross-section of extruded siliconeInner diameter (ID) of nozzle tipLayer height Speed of nozzle in the XY plane Shear rateIntroductionSilicone elastomers are inert materials unique for their wide range of useful properties, including thermal stability, bio-compatibility, hydrophobic nature, electrical insulation, low chemical reactivity and toxicity, and resistance to ultraviolet radiation [1].

These and other properties have distinguished silicone elastomers formany industrial applications including automotive, consumer, medical, and others. For example, in healthcare, the biocompatible and soft silicone elastomer make it ideal for maxillofacial prosthesis [2] and augmentation of soft tissue in plastic surgery [3]. Molding is currently the most common manufacturing process for silicone elastomers [4]. The mold, typically made of hardened tool steel, usually requires extensive capital cost and long lead times. Although molding is ideal for high volume mass production, the cost and lead time often limit the low volume production of custom silicone elastomer products. The ability for layer-by-layer additive manufacturing (AM) of silicone elastomer can overcome this barrier.

The ability for layer-by-layer additive manufacturing (AM) of silicone elastomer can overcome this barrier. In this study, the extrusion-based AM of moisture-cured silicone elastomers is explored to directly create the 3D freeform shape of silicone elastomer.Extrusion-based AM is a process whereby the viscous material is extruded through a nozzle for line-by- line building of a layer in an object [5]. Computer-controlled motion stages are used to control the movement of the nozzle for material deposition. There are two approaches for extrusion-based AM of silicone elastomer. The first is to use moisture-cured room temperature vulcanization (RTV) silicone elastomer, which is cured at the room temperature after extrusion from the nozzle. Another approach is to mix two types of silicone (commonly known as A and B) and then extrude the mixture out of the nozzle for AM.

Another approach is to mix two types of silicone (commonly known as A and B) and then extrude the mixture out of the nozzle for AM. The advantage of moisture-cured silicone elastomer AM is the simplicity and flexibility to fabricate soft freeform silicone elastomer parts. The soft silicone elastomer is targeted to have the elastic modulus of the work-material lower than 120 MPa. Most of current AM materials, including metals, ceramics, and plastics, have high elastic moduli to maintain the shape for layer-by-layer AM. When the extrusion-based AM technology is used for soft materials, many technical challenges arise. For example, the cross-sectional geometry, connection of lines and deformation of the soft silicone elastomer to achieve dimensionally accurate shapes are important topics of research.Extrusion-based AM of soft materials has been studied for fabrication of hydrogel scaffolds with specific shape and internal interconnected porous structure [6].

Tissue engineering is another application of soft material AM. Mironov et al. [7] developed a cell printer to stack soft materials (e.g. gels, single cells and cell aggregates) to build vascular and organ structures. The Freiburg Materials Research Center has developed an extrusion-based hydrogel and polymer AM machine for tissue engineering [8]. This approach has been applied for tissue scaffolds [9], biological matrices [10], fiber membrane [11], dental devices [12], and bio-composite scaffolds [13]. However, the research in extrusion-based silicone elastomer AM is still lacking. There is also a need for numerical modeling of the silicone extrusion process to predict the air pressure required for a specific nozzle size and flow rate, all critical parameters for this extrusion-based AM process.This research is aimed to narrow this gap by studying the extrusion-based AM of RTV moisture-cured silicone elastomer.

The experimental setup is presented and the process parameters are described in the following section. A numerical simulation is developed to model the size of free flowing silicone extrusion and predict the diameter and flow rate of the silicone. Results of experiments and simulation are compared and effects of key process parameters are discussed.Material and Experimental SetupSilicone materialThe silicone used in this study is the 890-SL RTV moisture-cured silicone (Dow Corning, Midland, Michigan). This is a one-part silicone elastomer that cures upon exposure to atmospheric moisture [14].Experimental setupThe experimental setup is shown in Fig. 1(a).

This is a one-part silicone elastomer that cures upon exposure to atmospheric moisture [14].Experimental setupThe experimental setup is shown in Fig. 1(a). The system consists of: (1) silicone syringe and the nozzle, (2) pressure regulator, (3) motion stages and (4) controller and data acquisition system.The silicone in a 25mm inner diameter syringe is extruded through a nozzle tip by controlling the air pressure inside the syringe using an electronic pressure regulator (Performus™ III by Nordson, Westlake, Ohio) until a steady-state flow rate is reached. The regulator has 3.5 kPa resolution to control the silicone flow rate.An ArduinoTM Mega 2560 controller is used to regulate air pressure to adjust the extrusion flow rate and synchronize with linear motion stages. Three linear stages (Model 200cri, Siskiyou, Grants Pass, Oregon) are used to guide the motion of the extrusion nozzle in X, Y, and Z directions (Fig.

2, there are five key input process parameters for the extrusion-based silicone AM process.Inner diameter of nozzle tip (di): This is the ID of the nozzle tip orifice, which determines the diameter and influence the flow rate of extruded silicone.Air pressure (P): The air pressure needs to match to the nozzle ID to achieve the target flow rate for silicone extrusion.Nozzle speed in the XY plane (v): The X and Y stages combine to control the nozzle speed in the XY plane to lay lines for silicone AM. The nozzle speed needs to match to the silicone flow rate.Layer height (t): The layer height is the distance in the Z direction from the nozzle tip to the surface of the previous layer, as shown in the layer deposition condition in Fig. 2(a).Distance between two adjacent silicone lines (c): The distance between two adjacent silicone lines, as shown in Fig. 2(c). This parameter affects the bonding and structural integrity of the silicone elastomer AM part.Fig.

2(c).The diameter d is measured using imaging processing technique. Fig. 3(a) shows the digital image of extruded silicone under the free flowing condition. Threshold intensity is used to process the data to either black or white pixels, as shown in Fig. 3(b). Along the Z direction, the number of black pixel is identified and plotted as in Fig. 3(c). The Z position of the nozzle tip OD (with diameter do) and the steady-state extruded silicone diameter (d) are identified and number of black pixels is denoted as Ntip and Nsilicone. The extruded silicone diameter d can be calculated using:dNsilicone Ntipd0The weight of extruding silicone under the free flow condition in a given time is measured using an electronic scale (Model HRB-103, Tree).

The weight of silicone is converted to volume to calculate the flow rate.(1)z0(a)2040n o i t c e r i d z6080100z(b)g n o l a120l e x i P140160180NsiliconeNtip200 20 22 24 26 28 30 32 34 36 38Number of black pixel(c)Fig. 3 Measurement of the diameter of extruded free-flowing silicone: (a) image of the camera with close-up lens, (b) pixel representation of the image, and (c) calibration of the nozzle OD and the measurement of d.The a and b of cross section of the silicone in the layer deposition condition are measured using image processing. A sample cross-section is illustrated in Fig. 4. The elliptical profile of the cross-section of an extruded silicone elastomer line is identified to estimate parameters a and b.Fig.4 Cross-section of extruded silicone lines.Experimental DesignFive sets of experiments are designed to investigate relationships between input and output parameters during the AM of silicone elastomer.Exp.

I – Effect of di: Under the free flow condition, the size di and the flow rate Q of the extruding silicone are affected by the nozzle orifice size di and air pressure P. In this study, four nozzles with di=0.41, 0.61, 0.84 and 1.19 mm are tested under P=207 kPa. The d and Q are measured.  Exp. II – Effect of P: With di=0.41 mm, the air pressure is set to be 138, 162, 207, 241, 276, 310, and 345kPa to measure the d and Q under the free flow condition.Exp. III – Effect of v: Effect of the nozzle speed matching with silicone extrusion speed is investigated. The nozzle speed v is controlled to be 15, 20, 25, 30, and 35 mm/s to study the shape of silicone lines with t =c = 0.84 mm, di=0.41 mm, and P =262 kPa.Exp. IV – Effect of t: Experiments with t = 0.5. 0.6, 0.7, and 0.84 mm are conducted with di= 0.41 mm, P =262 kPa, v = 25 mm/s, c = 0.84 mm (the diameter of free flowing silicone).Exp.

V – Effect of c: Experiments with c= 0.5, 0.6, 0.7, and 0.84 mm are conducted with di= 0.41 mm, P =262 kPa, v = 25 mm/s, and t = 0.84 mm.Numerical Modeling of Silicone FlowThe COMSOLTM Multiphysics (v4.4) is used to simulate the silicone extrusion process.Governing equationsThe silicone is modelled as an incompressible fluid with the mass conservation: u0where u is the velocity vector of silicone.The dynamic equilibrium of the silicone can be expressed by the momentum equation.u t where ρ is the density, μ is the viscosity, and p is the hydrostatic pressure.） u -（ Tup((u+ u))+0The non-Newtonian Carreau model is used to describe the viscosity of silicone based on the viscosity ofthe silicone at infinite shear rate μ∞ at zero shear rate μ0.n 21（） [1 (2 ) ] 0 -  (4) where λ and n are material constants.

Based on the material properties provided by the material supplier (Dow Corning), these parameters are as follows: μ∞=0, μ0=62.5 Pa.s, λ=1.73x10-2s, n=0.5, and ρ=1039 kg/m3.A laminar two-phase flow model [15] is used to simulate the silicone flow. For tracking the interface between air and silicone, the level set method [16] is applied. The level set method use a function φ to describe the interface between the silicone and air. The value of φ varies from 0 to 1 in a transition layer near the interface. On the interface, the value of φ equals to 0.5. The convection of the level set function and interface between two fluids is [16]. t )))] 0( (1 [()u( || where ε is proportional to the thickness of the convection layer and γ determines the amount of the reinitialization [16].

Meanwhile, the level set function can describe the smooth change of density ρ and viscosity μ across the interface.=  =  airair((silicoensilicoen)   air )   airMesh and initial and boundary conditions(2)(3)(5)(6)An axisymmetric model, as shown in Fig. 5, is used to represent the silicone inside the syringe chamber and the nozzle. An adaptive meshing operation is applied to model the silicone extrusion process, as illustrated in the close-up view around the nozzle tip. Boundary conditions of the silicone and air domains at stationary (zero velocity condition) are defined in Fig. 5. There are three boundary conditions, denoted as BC I, BC II and BC III. BC I is the inlet of the silicone, which is maintained at the pressure P, at the top of the syringe chamber. BC II is the no-slip condition on the syringe wall between the silicone and the control volume boundary of air.

BC II is the no-slip condition on the syringe wall between the silicone and the control volume boundary of air. BC III is the outlet at the bottom of the air control volume with 101.3 kPa. The interface between the silicone and air at the start of the simulation is assumed at the nozzle tip.Fig. 5 Boundary conditions of the axisymmetric model at the start of simulation.Fig. 6 The value of φ of level set function for free flowing silicone.Results and discussionsResults of EXP. IFig. 6 is the value of φ for silicone extruded through the nozzle in steady-state free flowing condition. The interface between the silicone and air is determine with φ=0.5. The diameter, d, of the extruded silicone is slightly expanded after extrusion from the nozzle orifice. Due to the gravity effect, the diameter of silicone is reduced. The largest diameter of the silicone out of the nozzle is identified and used for parameter d, presented in Fig. 7.Fig.

Due to the gravity effect, the diameter of silicone is reduced. The largest diameter of the silicone out of the nozzle is identified and used for parameter d, presented in Fig. 7.Fig. 7 shows good agreement between the experimentally measured and numerically predicted d under freeflowing condition in Exp. I with di=0.41, 0.61, 0.84 and 1.19 mm. The d is about 10 to 15% larger than di.The model predicted and experimental measured Q also has a good agreement. Higher di corresponds to alarger area of the nozzle orifice and the quadratic increase in the Q.Experiment d Simulation d1.2401.135301.0250.90.820150.7100.650.500.40.40.60.81.01.20.40.60.8 Nozzle size di（mm）1.0Fig. 7 Experimental and modeling results of d and Q vs.di in Exp. I.Results of EXP. IIFig.

7 Experimental and modeling results of d and Q vs.di in Exp. I.Results of EXP. IIFig. 8 shows the inverse relationship between P and d, which is validated by both experimental measurement and numerical simulation for P = 138, 162, 207, 241, 276, 310, and 345 kPa. Under low P (138kPa), the silicone expands and the diameter d is 0.55 mm, much larger than the 0.41 mm di. The d is reduced to slightly lower than di at the highest pressure (345 kPa). As expected, the Q increases at a higher P, as shown in Fig. 8.Experiment d Simulation d0.5880.5670.540.5260.500.480.46) s /3m m(540.44Q30.4220.400.3811001502002503003500 100150200250300350Fig.8 Experimental and modeling results of d and Q vs. P in Exp. II.1.2The agreement between experiment measured and model predicted d and Q in Fig. 8 is also reasonable.Good agreement between the experimental and modelling results in Figs.

P in Exp. II.1.2The agreement between experiment measured and model predicted d and Q in Fig. 8 is also reasonable.Good agreement between the experimental and modelling results in Figs. 7 and 8 validates the approach using level set function for modelling of the silicone flow for extrusion-based AM.Results of Exp. IIIThe picture of the silicone lines for v = 15, 20, 25, 30, and 35 mm/s in Exp. III is shown in Fig. 9. Results of the long and short radii of the cross section are provided in Table 1. Under the given Q = 3.6 mm3/s, at the highest v (35 mm/s), the silicone line is discontinuous. At a lower v, the silicone line becomes wider. The silicone elastomer can be stretched or compressed to a certain range.

At a lower v, the silicone line becomes wider. The silicone elastomer can be stretched or compressed to a certain range. Under a given extrusion flow rate, v can vary within a range for extrusion-based AM.Table 1 Cross section with different nozzle speedNozzle speed v (mm/s) Long radius a (mm) Short radius b (mm)151.0150.425200.6450.425250.420.42300.360.415350.2650.405Cross sectionContinuous Very wideContinuous WideContinuous NormalContinuous narrowCrack Very narrowFig. 9 The silicone elastomer line vs. nozzle speed v in Exp.III.Results of Exp. IVThe a, b, and the cross-section views of AM silicone elastomer lines for t = 0.5. 0.6, 0.7, to 0.84 mm in Exp. IV are shown in Fig. 10. As expected, a lower t compresses the silicone after extrusion, increases a, and decreases b. The b is about half of the t due to compression of the silicone between the nozzle and the layer surface.

As expected, a lower t compresses the silicone after extrusion, increases a, and decreases b. The b is about half of the t due to compression of the silicone between the nozzle and the layer surface. The layer height t influences the silicone elastomer AM accuracy and efficiency. If t is large, the silicone elastomer AM is more efficient as the number of layers decreases, while the accuracy of prints is compromised. A balance is required in terms of t. As t increases while c remains the same, the cross section turns from an ellipse to a circle and a becomes smaller.Long radius(a) Short radius(b)0.500.440.420.480.400.380.360.460.340.320.440.300.280.420.260.240.450.500.550.600.650.700.750.800.85Fig. 10 Cross section vs. layer height tResults of Exp. VThe a, b, and the cross-section view of AM silicone elastomer lines for c = 0.5, 0.6, 0.7, 0.84 mm in Exp. V are shown in Fig. 11. The b is about 0.42 mm, half of the t.

VThe a, b, and the cross-section view of AM silicone elastomer lines for c = 0.5, 0.6, 0.7, 0.84 mm in Exp. V are shown in Fig. 11. The b is about 0.42 mm, half of the t. Under a narrow (low) c, the a is lower due to the squeezing between lines. The c affects the bonding strength of silicone lines and structural integrity of AM silicone elastomer parts. A larger c could be used for internal structure, while a narrower c may be better for surfaces.Long radius(a) Short radius(b)0.481.41.21.00.460.80.60.40.440.20.00.20.420.40.60.450.500.550.600.650.700.750.800.85Fig. 11 Cross section vs. distance between silicone lines cConclusionsThis research demonstrated the feasibility of extrusion-based AM of silicone elastomers and systematically investigated effects of five input parameters on the shape and flow rate under both free flowing and layer deposition conditions.

The ComsolTM multiphysics modelling with level function to predict the silicone-air interface had also proven to be feasible to simulate the silicone flow during the extrusion process. This paper presented some initial but fundamental and quantitative investigations on the silicone elastomer AM process. The mechanical properties of the AM silicone elastomer parts will be further studied in our future study.AcknowledgementsWe acknowledge the support from Dow Corning.References[1] Moretto H, Schulze M, Wagner G, Silicones BE, Hawkins S, Russey W, Schulz G (1993) Ullmann'sEncyclopedia of Industrial Chemistry, in, Verlag Chemie: Weinheim, Germany.[2] França DCC, De Castro AL, Soubhia AMP, Tucci R, Goiato MC. Biocompatibility evaluation of 3 facialsilicone elastomers. Journal of Craniofacial Surgery; 2011. 22(3): 837-840.[3] Chasan PE. The history of injectable silicone fluids for soft-tissue augmentation. Plastic and reconstructivesurgery; 2007.

b@Home Model 2: Towards Ubiquitous Personal Fabrication Devices Jeffrey I Lipton1, Daniel Cohen2, Michael Heinz1, Maxim Lobovsky1, Warren Parad3, Garrett Bernstein1,4, Tianyou Li4, Justin Quartiere4, Kamaal Washington2, Abdul-Aziz Umaru2 Rian Masanoff5, Justin Granstein6, Jordan Whitney4, Hod Lipson2,4 1 Applied and Engineering Physics 2 Mechanical and Aerospace Engineering 3 Electrical and Computer Engineering 4 Computer Science 5 Department of Government 6 Department of History Cornell University, Ithaca, NY 14850 USAAbstract: The open-architecture, open-source Fab@Home platform has proven to be an important system within the SFF community. In order to facilitate wider spread of the Fab@Home platform and SFF throughout the world, we aimed to improve critical aspects of the system, and business model. By changing the electronics package and streamlining the mechanics, the cost of the system was brought from $2500 to $1600.

By changing the electronics package and streamlining the mechanics, the cost of the system was brought from $2500 to $1600. By changing the business model we hope to transform the SFF market and spur innovation.Introduction:Solid Freeform Fabrication (SFF) has the potential to transform the world. Personal Fabrication devices, “fabbers”, can enable the individual to manufacture complex objects on demand. Fabbers change the methods used to design, make, deliver, and use products. They eliminate the barrier to production that often stifles personal innovation. Rather than needing larger capital investment, or extensive skill sets, the end user can immediately transmute a CAD representation into a functional physical object. The fabber revolution could be as important, and transformative, as the digital revolution by allowing for the “democratization of invention”.

The fabber revolution could be as important, and transformative, as the digital revolution by allowing for the “democratization of invention”. (1) The goal of the Fab@Home project is to spark innovation and development of personal SFF systems.The majority of SFF systems on the market today are large, expensive proprietary machines. Few are below the $20,000 mark. These high cost machines limit the spread of SFF technologies by erecting a barrier to personal ownership. They stifle innovation, by making systems technologies proprietary. The business model of most printers is similar to that of the old mainframe computers from before the digital revolution. They use a vertical business model, where a single company makes all material, hardware and software for the system. Fab@Home seeks to transform all of this, by establishing a cheap open-source platform.

Fab@Home seeks to transform all of this, by establishing a cheap open-source platform. Rather than using a vertically integrated business model, the Fab@Home project seeks to transition to a horizontal model, similar to the modern PC era, where different companies, and make the materials, hardware, and software. (2)Background:The objective of the model 1 Fab@Home was to lower the cost of involvement in SFF, and encourage others to develop and invent new technologies. To this end the model 1 has been extremely successful. The cost of a Model 1 printer brought the cost of ownership to $2300 for parts. This price reduction has made it possible for individuals, schools, and startups to purchase their own printers. Printers have been deployed across several continents. (3) Several companies have formed around the Fab@Home idea and begun to manufacture and distribute them. Additionally the model 1 has helped to spark interest in the general public about SFF technologies.

Additionally the model 1 has helped to spark interest in the general public about SFF technologies. It attempted to make a printing system similar to the standard business model of printers; the hardware and software of the Fab@Home were integrated into a single platform.Several limitations of the Fab@Home printer platform have presented themselves over its life. Though the printer is significantly less expensive than most commercial printers, the price mark is still a barrier for many people who are interested in the project. Since it is an open- source community backed program, a high price tag is prohibitive since there is no one to return a kit too if one is dissatisfied; trust and faith are required. Additionally the difficulty of assembly and modification has frustrated many users and hurt the projects dissemination. The community has been poorly lead, and since its inception, has not grown significantly.

The community has been poorly lead, and since its inception, has not grown significantly. There has been no direction or guidance, only limited support.The Reprap project has been extremely successful in cultivating an open source community. It has given direction, guidance, and updates to the community on a regular basis. As a result they have an active user base and community. While its high skill level requirement tends to limit the community to more technically inclined users, these users are the ones who develop and further open-source projects. Like all other printers on the market, it is based on a vertical business model. (4)Methods: HardwareIn order to improve the quality and cost of the Fab@Home system, new hardware needed to be developed. The largest cost in the Fab@Home Model 1 kit were the transmission, at $662.52, and electronics with motors ($367.78+ $610.59) (3). These systems were responsible of the limitations in quality, speed, and complexity of the Fab@Home.

These systems were responsible of the limitations in quality, speed, and complexity of the Fab@Home. Complete redesign of the transmission and motors necessitated the redesign of the entire model 1 chassis system.The Motors and electronics of the model 1 limit the system to low speeds. At higher speeds, the system begins to drift as a result of the stepper motors loosing steps. This error creates a print speed cap at 10 mm/s. There is no error correction in the system, so all errors compound. This causes long builds to be difficult. The drivers associated with this system are not perfectly stable; often the drivers will crash and require the complete restart of the host computer. The assembly of this system, ad its exposed design often intimidates would be users.The electronics and motors were replaced with an integrated single supplier commercial system. The Snap Motor system combined the controls, amplifier, motor and optical encoder into a single unit.

The Snap Motor system combined the controls, amplifier, motor and optical encoder into a single unit. (5)The Cabling was accomplished by use of standard cat5 cables. The systems commercial drivers ensured support for the platform’s electronics.In order to improve the transmission, the system was switch from lead screws to timing belts, and the parts were homogenized. On the Model 1, each axis used different linear bearings , and different shafts. This increased the complexity and cost of the kit. In order to reduce complexity, the system was designed to use identical shafts and bearings for each axis. Transmission was accomplished using timing belts, pulleys and idlers. This system inherently costs less and eliminates the need to purchase an interface between the transmission and the chassis, rather that parts are made entirely from acrylic. Additionally all but one, non critical component of the system was switched to metric.

Additionally all but one, non critical component of the system was switched to metric. This enabled a more global appeal to the model 1 since outside of the United States, standard parts are difficult to acquire.New tool heads were designed to give the printer more versatility. Aside from a traditional syringe based displacement tool. A pressure and valve based tool was developed. A pressure source forces material from the syringe reservoir at a fixed rate while the valve is open. Additionally a Dremel mount was developed for the printer.Methods: SoftwareRedesigning the software was critical to forwarding a horizontal business model. The old system had a fully integrated print planning and printer control program. A change to any aspect of the system required extensive rewriting of the software. We sought to enable the interoperability and independent development of several layers of programming. This was accomplished by dividing the software as seen in figure 1.

We sought to enable the interoperability and independent development of several layers of programming. This was accomplished by dividing the software as seen in figure 1. In order to test this model, the FabStudio, .Fab files, and FabInterpreter were created.A “.fab” file type was designed as an XML based file which allows for the efficient transfer of commands and information. XML was selected because of its hierarchical nature and widespread usage. A “.fab” file contains relevant information about the tool, material settings, and paths in the form of tagged information. It was designed to allow future interpreters and planners to add commands and sequences that have not been predicted. XML’s nature allows for parsers to ignore unknown commands. These commands may include flow control such as loops, or logical structures such as “if” statements.The API provided with The Snap Motors was used temporarily to build the interpreter (5).

These commands may include flow control such as loops, or logical structures such as “if” statements.The API provided with The Snap Motors was used temporarily to build the interpreter (5). This allowed for the rapid development of the initial control hardware. The interpreter allows for the operation of the Model 2 printer and can convert the information in a .fab file into commands to the printer with result in physical objects. The client program, named FabStudio, uses path planning algorithms to convert an STL, tool and material information into a .fab file.

The client program, named FabStudio, uses path planning algorithms to convert an STL, tool and material information into a .fab file. It was created by modifying the model 1 program and removing all hardware interactions.PlannersCommand MediumPrinter ControlsControl structureFabStudioFoodFab.fab FilesFab InterpreterBioPlannerMasterCamG-code filesG Code interpreterAPIScanner Control(a)Control StructureAPIHardware interfaceSnap Motor DriversModel 1 DriversFuture Model 3Reprap DriversNextFabHardwareModel 2valve toolone syringetwo syringeFuture Model 3DarwinNextfab(b) Figure 1: The new printer control scheme(a) could enable a transition to a horizontal business model for SFF devices. If SFF systems controls could conform to a set API structure, it would be possible to separate the hardware and software entirely.

If SFF systems controls could conform to a set API structure, it would be possible to separate the hardware and software entirely. Just as AMD and Intel processors do not require different operating systems, SFF software would operate any number of printers and tool heads (b).Methods: Community DevelopmentIn order to encourage community growth and interest in the project, a new website was designed. The old wiki based website’s visual appeal was lacking. It did not immediately draw a new user into the project. The information on the site was disorganized and often unhelpful. The new website was designed to educate and entice new users, while organizing information in useful ways for established users.The site was divided into two main parts. The front end of the site was made as a static page powered by Drupal. (6) It was filled information and visuals that encourage visitors to become users.

The front end of the site was made as a static page powered by Drupal. (6) It was filled information and visuals that encourage visitors to become users. It was designed to be a manifestation of the guidance and leadership in the project that was lacking. The Back end of the site was made from a wiki with integrated blog and forum. Previously the wiki and forum were separate. The goal of integrating them into a single site was to encourage a sense of organization and community. The blogging was added to the website in order to encourage users to return to the site frequently, and to consolidate information about the end users activities.Results: HardwareThe design of the model 2 reduced the price and build-complexity of the Fab@Home system, while increasing its aesthetic appeal. figures 3 shows various parts of the Model 2. The positioning of the motors and elimination of custom wiring, along with the reduced visual impact of the transmission gave the Model 2 a cleaner design.

The positioning of the motors and elimination of custom wiring, along with the reduced visual impact of the transmission gave the Model 2 a cleaner design. Table 1 shows the components and cost of the Model 2. The modification of the electronics and transmission greatly reduced the cost and number of parts.The skills required to build a Fab@Home have been greatly reduced. In order to build a Model 2 from a kit, one needs only a hex wrench set and a soldering iron for thermoplastic inserts. One need not program a microcontroller, or strip and solder any wires since the Snap Motor system is preprogrammed and uses cat 5 cables.The optical feedback eliminated drift provided that the transmission is properly tightened. A single encoder count corresponded to 0.006 mm of travel in the x direction, 0.0035 mm in y, and 0.0036mm in z.

A single encoder count corresponded to 0.006 mm of travel in the x direction, 0.0035 mm in y, and 0.0036mm in z. The error correction built into the motors allowed for the system to automatically return the system to its proper position after an error was induced by exerting a force on the deposition tool during a print.The printer has been successfully used with a top movement speed of 100mm/s with an acceleration of 100mm/s2. This is a significant improvement over the maximum speed of the model 1 at 25 mm/s theoretically, and 10mm/s in practice. Objects have been printed on the Model 2 at twice the build speed of the Model 1. The speed ceiling of the Model 2 appears to be at 20mm/s during a print, though with further refinement it may go higher.(a)(b)(c)(d)(e)(f)Figure 2: The model 2 Fab@Home (a) comes with a two syringe displacement based deposition tool (c) , and uses the same shafts and bearings(d).

The planner has created dozens of .fab files. These files were stored to allow for instant access by the interpreter. A standard text editor can be used to edit the paths and print parameters generated during planning. This has been used as a mean of error correction. Occasionally the planner will add unnecessary paths, or paths that are too small, and manual editing can correct planning errors. The interpreter was used to convert .fab files into objects.(a)(b)(c) Figure 3: The Fab@Home control software has been divided into a two independent pieces of software, the planner, FabStudio (c), and the controller call FabInterpreter (a). These communicate via .fab files (b).Results: Community DevelopmentThe new site successfully provides a visually appealing lure which draws users to the project. Figure 5 shows the sitemap of the website. All front pages are used to connect visitors with the relevant information on the wiki, and discussions on the forum.

Figure 5 shows the sitemap of the website. All front pages are used to connect visitors with the relevant information on the wiki, and discussions on the forum. The site has not gone active yet, and is still under construction.Figure 4: The sitemap for the Fab@Home community was designed to attract new users and guide the project. Arrows represent links.Discussion:The transition to a horizontal business model for SFF devices began of the Fab@Home systemwith model 1. The syringe based deposition tool used on the model 1 allowed any material to be extruded. Kraftmark is an independent materials supplier who invented two different types of materials for 3d printers. (7) One of which, FabUnotm is provided in the standardized syringe container. Had Kraftmark sought to innovate on printing materials prior to the model 1, they would have need to invent and entire printing platform.

Had Kraftmark sought to innovate on printing materials prior to the model 1, they would have need to invent and entire printing platform. The separation and standardization of the Fab@Home software allows each aspect of the software to be developed independently, provided it follows interaction standards.The FabStudio and FabInterpreter programs demonstrate the ability to separate the developmentof printer controls from print planning. The Fab@Home has been used in biological experiments, cake decorating and SFF. Each of these uses could benefit from different planning programs, algorithms and commands, but could use the same control software. Future standardization of software hardware interactions would enable interoperability of printers.

Future standardization of software hardware interactions would enable interoperability of printers. Ideally corporations, individuals, and other open source groups would develop the separate aspects of the printing process.While the new software and business model will remove the barriers to innovation once involved in SFF, the Model 2 printer will remove the barriers to involvement. The key barriers to involvement are price, skill level, and belief in usefulness. The price reduction from the chassis and transmission redesign will enable more individuals and corporations to invest in SFF technologies. The newly designed motor and electronics system reduces the skill level needed. The Dremel attachment will ensure a minimal usefulness level to the system. Engravers and CNC mills are a well established field. People understand their usefulness. By incorporating this functionality into the device, people will have confidence to invest in a printer for personal use.

People understand their usefulness. By incorporating this functionality into the device, people will have confidence to invest in a printer for personal use. A strong community will give people confidence in the longevity and support for their investment in SFF devices.Future Work:Working towards personal fabrication devices will demand a great deal of effort. The next step for the Fab@Home project is to develop a support infrastructure based in Cornell University and around the world. An undergraduate project team should be created in order to provide support and constant innovation to the Fab@Home project. Using the resources of a project team and a resuscitated community new tools and software, and hardware should be developed. This will include a filament based deposition tool to enable plastic printing similar to Reprap and other low cost printers.

This will include a filament based deposition tool to enable plastic printing similar to Reprap and other low cost printers. Additionally, higher capacity reservoirs based around a valve powered tool to encourage extended use of a printer to create large objects. A custom electronics and motor option should be created. While this would deviate from the Fab@Home projects use of exclusively off the shelf parts, it would help reduce the entry cost for experienced users and commercially deployed machines. A new independent API standard should be developed in collaboration with other open-source printers to encourage standardization and stratification of SFF device development.Conclusion:The spread and advancement of personal fabrication technologies requires that barriers to ownership and innovation be dramatically degraded. In order to accomplish this, a new printer, and business models, and revitalized community were needed.

In order to accomplish this, a new printer, and business models, and revitalized community were needed. The new printer design greatly reduces the price and skills level barriers while improving technical performance. The new website will help translate interest into participation in the community. By standardizing hardware and software interactions, innovations in any aspect of printing technology can be instantly transferred across all printer platforms. Together these advancements forward us towards ubiquitous personal fabrication.Acknowledgements:This research was supported in part by NSF Creative-IT grant 0757478 and the DefenseAdvanced Research Projects Agency (DARPA), Defense Sciences Office (DSO) under the Programmable Matter program.Works Cited 1. Hippel, Eric Von. Democratizing Innovation. Cambridge, Mass. : MIT Press, 2005.2. Ranjay Gulati, Mohanbir S. Sawhney, Anthony Paoni, J.L. Kellogg on technology & innovation. s.l. : Jon Wiley & Sons, 2002.3.

B@HOME MODEL 3: A MORE ROBUST, COST EFFECTIVE AND ACCESSIBLE OPEN HARDWARE FABRICATION PLATFORMJeffrey Lipton. Robert MacCurdy, Matt Boban, Nick Chartrain, Lawrence Withers III, Natasha Gangjee, Alex Nagai, Jeremy Cohen, Karina Sobhani Jimmy Liu, Hana Qudsi, Jonathan Kaufman, Sima Mitra, Aldo Garcia, Anthony McNicoll, Hod Lipson Cornell University, Creative Machines Lab, Ithaca NY 15850Abstract Solid Freeform Fabrication is transitioning from an industrial process and research endeavor towards a ubiquitous technology in the lives of every designer and innovator. In order to speed this transition Fab@Home Model 3 was created with the goal of expanding the user base of SFF technology by lowering the skill and price barriers to entry while enabling technology developers to leverage their core competencies more efficiently.

The result is a device, which is modular with respect to tool heads, fabrication processes, and electronics controls, costs under $1000, and requires only a simple tool set to assemble.IntroductionOver the last few years, we have seen an explosion in the personal 3D printer market. Companies in The United States, Europe, China and elsewhere are offering new options to consumers. Many of these devices are descendants of the early open hardware 3D printers. Most of these new units are being offered as cheap DIY kits that allow the user to assemble their own machine. The overwhelming majority of personal 3D printers use FDM manufacturing to make plastic objects; few attempt any other material, process or functionality natively. This limits the scope of potential users. While some early adopters and technology enthusiasts have a need to rapidly fabricate simple plastic components, few other potential users of personal 3D printers have the same need.

While some early adopters and technology enthusiasts have a need to rapidly fabricate simple plastic components, few other potential users of personal 3D printers have the same need. To them it remains a curiosity rather than a useful tool.In order to expand the user base of personal 3D printing technology, it is necessary to transform single material 3D printers into general-purpose personal manufacturing devices. Such personal fabricators would leverage the infrastructure of 3D printers to accomplish more traditional manufacturing processes. By combining additive and traditional manufacturing capabilities on the same machine, it is possible to dramatically expand the potential user base.Further limiting the adoption of 3D printing by new user groups is the cost, technical skill requirements, and ease of access of a printer. Most of the kits on the market require substantial technical skill to assemble and operate.

Most of the kits on the market require substantial technical skill to assemble and operate. A natural solution to this problem may be to offer fully assembled units; however, the complexity and technical requirements of the printer designs leads to extremely high labor costs, often putting the assembled units out of the price range of most potential users. (Makerbot Industries)Most of the current generation of personal 3D printers are in the vein of the first Model 1 Fab@Home units and traditional commercial machines. They are designed as a single integrated system, requiring the developer to develop novel electronics, chassis, tool heads and control software. (Makerbot Industries) (3D Systems) (pp3dp) This limits the speed of development andforces manufacturers to be well versed in mechanical, electrical, and software engineering. The diversity of skills required will inevitably lead to higher development costs and therefore higher product costs.

The diversity of skills required will inevitably lead to higher development costs and therefore higher product costs. In order to advance the transition to SFF ubiquity the Model 3 Fab@Home system must once again lower the user’s barriers to entry, increase the ease of modification, and expand the usage scope of the machine.BackgroundThe Fab@Home Project began as an open hardware project in 2005 under Hod Lipson and then doctoral student Evan Malone. The Model 1 Fab@Home was introduced in 2007 as one of the first low cost 3D printers. (Evan Malone, 2007) It was similar to traditional SFF devices in that it was an integrated system with the path planning and motion control software, electronics, chassis and tool head all tightly integrated. Unlike other 3D printers, it was capable of using a wide variety of materials. In 2009, the Fab@Home Project released the Model 2 3D printer design.

Unlike other 3D printers, it was capable of using a wide variety of materials. In 2009, the Fab@Home Project released the Model 2 3D printer design. It focused on lowering the barrier of access by dropping the cost, and was the first system to decouple the tool head from the chassis to ease the modification process. (Lipton, et al., 2009) Several hundred Model 1 and Model 2 units have been by users. These were self-reported by users on the Fab@Home website.Open Hardware is a nascent concept akin to the open software movement, which began in the 1980’s. In 2011, an international community of developers accepted the first definition of open hardware. According to the first definition,Open source hardware is hardware whose design is made publicly available so that anyone can study, modify, distribute, make, and sell the design or hardware based on that design. The hardware's source, the design from which it is made, is available in the preferred format for making modifications to it.

The hardware's source, the design from which it is made, is available in the preferred format for making modifications to it. Ideally, open source hardware uses readily-available components and materials, standard processes, open infrastructure, unrestricted content, and open-source design tools to maximize the ability of individuals to make and use hardware. Open source hardware gives people the freedom to control their technology while sharing knowledge and encouraging commerce through the open exchange of designs. (Open Hardware Summit)Fab@Home has been operating under this model since 2007. The designs are released under the BSD license and rely on off the shelf parts, and manufacturing processes which are widely available for individual usage, specifically laser cutting.The Fab@Home project is currently working on the Fab@School project.

The Fab@School project aims to revolutionize STEM education by introducing engineering into the classroom to integrate science technology and math education into a tangible activity. The project uses digital fabrication technology to enable students and teachers to make abstract notions in mathematics tangible. The program will begin in the second grade and build up towards high school.For a personal fabrication device to be successfully integrated into an elementary school classroom it must meet strict safety concerns and must be accepted by the teachers. The Curry School of Education in UVA uses a principle of least change. In order to ensure that teachers will migrate to a new method, it must build a bridge from current methodologies. Most classrooms use Elision die cutters and other forms of 2D paper fabrication.

In order to ensure that teachers will migrate to a new method, it must build a bridge from current methodologies. Most classrooms use Elision die cutters and other forms of 2D paper fabrication. Therefore, the potential user base of elementary school students and teachers must have a bridge built between their current usage of manufacturing technology and 3D fabrication as described in the introduction.Methods: Barrier to EntryThe Model 3 needed to reduce the difficulty of assembling the unit by reducing the skills required and by reducing the time required to build a unit. A universally accepted standard for low skill assembly is IKEA. IKEA produces items, which often require only a hex wrench set to assemble. The Model 2 Fab@Home required a soldering iron to be used install thermoplastic inserts. This proved to be intimidating to many non-technical people.

The Model 2 Fab@Home required a soldering iron to be used install thermoplastic inserts. This proved to be intimidating to many non-technical people. Therefore, the Model 3 needed to eliminate the thermoplastic inserts in order to ensure the maximal amount of people would be comfortable with attempting assembly of a unit.In order to reduce the build time of a Fab@Home, the system needed to be made as simple as possible. This was accomplished by having fewer screws and interlocks. Screws and interlocks on the Model 2 have been a source of frustration and failure. These spots often are the sources of cracking in acrylic. Variation in the thickness of the acrylic would often require the users of the Model 2 to file the pieces to ensure a proper fit.The system was designed to reduce the amount of laser cutting required, use more cost effective parts, and electronics other than the JrKerr SnapMotor system were adopted in order to lower costs of ownership.

Laser cutting time was reduced by attempting to minimize the perimeter of all parts. This leads to geometrically simpler parts. The reduction in the number of screws used greatly aided this process by eliminating the holes for the screws and nuts from the interior and exterior perimeter of many parts. Laser cutting times were recorded by cutting the parts out of 6mm acrylic on a 35-watt Epilog Legend laser-cutter at 3% of maximum speed and at full power.With the Model 2 Fab@Home system, ease of access was a major barrier to user adoption. Often those community members providing kitting services were unable to stock the necessary parts from the vendors. Users in countries other than the United States often had problems ordering the parts from American vendors, and could not find suitable replacements in their own countries.

Users in countries other than the United States often had problems ordering the parts from American vendors, and could not find suitable replacements in their own countries. Since the Model 2 had been designed around specific parts from specific vendors this often created delays for users who wanted to adapt the system design to locally available parts. Therefore, the system needed a more flexible chassis design, which would allow for a more flexible supply chain. By under-constraining the dimensions of parts, it is possible to make a system, which will be more adaptable.Methods: Ease of ModificationThe Model 2 Fab@Home successfully decoupled the printer chassis from the tool heads. This lead to a wide variety of tool heads on the Model 2 system. Unfortunately, the space of possible tool heads was limited by the tight integration of the JrKerr SnapMotor system into the chassis of the Model 2. This necessitated the use of SnapMotors for the tool heads.

This necessitated the use of SnapMotors for the tool heads. These motors drove up the cost and limited innovation. For Model 3 it was necessary to separate the selection of motors and electronics control systems from the chassis. Other 3D printers have had interchangeable controller boards; however, they always required use of specified motors. In RepRap designs, the NEMA stepper motor standards were always specified in the chassis design. However, the electronics were easily changed. In order to ensure that the system could work with as wide a range of motor as possible, the Model 3 system uses interchangeable motor bays.These bays allow for the mounting of various motors onto the chassis without the need for extensive modification of the machine’s design.Methods: Expanded functionalityFab@Home systems are, in essence, a 3-axis robotic gantry system with interchangeable tool heads.

The robotic nature of the Fab@Home naturally leverages itself to other forms of manufacturing aside from deposition. With the Model 2, the system was extended into CNC fabrication using a Dremel flex-shaft mount tool head. In order to build the bridge between potential users’ current needs and Solid freeform fabrication system, the Fab@Home’s set of tools was expanded to included: vinyl cutting, pen plotting, foam cutting, pick and place assembly, microscopy, automated pipetting, 8 material printing, and plastic printing. These tools were designed to use a variety of different motors and leverage the interchangeable mounting system developed in the Model 2 and the motor selection flexibility of the Model 3For the elementary school environment, one needs to add additional safety features, which are not typically used in a hobbyist or research environment.

In order to mitigate the smells from the fabrication process and prevent users from putting their body parts in the machine during operation, a safety hood was developed which can enclose the Model 3. This allows users to customize the safety level of their machine. In the event that a user does obstruct the machine’s motion with their body, the system can be equipped with tension-limiting devices that ensure the system can apply a maximum force on the user below the point of causing bodily harm.Results: Barrier to EntryFigure 1: The Model 3 Fab@Home with JrKerr SnapMotors and a Single syringe displacement toolThe techniques employed greatly reduced the assembly difficulty of the Model 3 relative to the Model 2. The Model 3 requires hex wrenches, Philips screwdrivers, scissors and a c-clamp to assemble. These are all tools that almost any individual would feel comfortable using.

The Model 3 requires hex wrenches, Philips screwdrivers, scissors and a c-clamp to assemble. These are all tools that almost any individual would feel comfortable using. Belt attachment mechanisms were decoupled from their assemblies to allow for easier adjustment and installation of the belt drives. The system was designed to be made in distinct sub-assemblies, which could be made in parallel and then integrated together.The simplification of the design allowed an individual to build a prototype of the Model 3 using the JrKerr SnapMotors in a little over one hour of work. This individual was the designer of the Model 2 and Model 3 Fab@Home systems. He typically required 6-8 hours to assemble a Model 2 fabricator. Many users reported needing approximately 20 person-hours to assemble a Model 2 unit, 2.5 times as long as it took the designer. Therefore, it should take a naive user about 3 hours to assemble a Model 3, provided they had all the relevant documentation.

Therefore, it should take a naive user about 3 hours to assemble a Model 3, provided they had all the relevant documentation. This is a dramatic improvement over the average two-day build process of other machines.By reducing the perimeter of the parts and replacing the bottom plate of the Model 2 with a multiple plate system , it was possible to reduce the cutting time to 298 minutes on the Epilog Legend, without using edge doubling. The Model 2 often took upwards of 400 minutes to lasercut on the same machine. This should help reduce the cost of laser cutting for users who do not have access of their own laser cutter.The Model 3 has significantly lower transmission and chassis costs than the Model 2. By changing more pulleys to plastic pulleys, switching to 8mm shafts, and using sleeve bearings rather than ball bearings, the Model 3 transmission costs were reduced to $263 from $399 for the Model 2 a 34% cost reduction for the same functionality.

The removal of inserts and the adjustable base pads allowed the Model 3 chassis to cost $179.53 compared to $221.97 for the Model 2, a savings of 19%. (Lipton, et al., 2009)The most substantial savings come from the interchangeable nature of the motors. This allows the system to use a variety of different more cost-effective motion systems. While the JrKerr SnapMotors are the simplest to install for a novice user, they are expensive. Other motion systems, including the new hobby servo based electronics developed by the Fab@Home Team, are half as expensive for a 3 axis motion system, and nearly a third as expensive for a 5 motor system(see figure 2). In total, a Model 3 using a dual-syringe displacement tool and the servo electronics should cost $935 plus laser cutting and shipping costs. A Model 2 with a dual-syringe displacement tool costs $1600 plus laser cutting and shipping.With the Model 2, the metal pulleys from SDP/SI often caused supply chain issues.

A Model 2 with a dual-syringe displacement tool costs $1600 plus laser cutting and shipping.With the Model 2, the metal pulleys from SDP/SI often caused supply chain issues. The design of the Model 3 idler pulley system allows the idlers to vary by upwards of 5mm in length and 3mm in diameter. This will allow users to adapt the system to the locally available parts.

The design of the Model 3 idler pulley system allows the idlers to vary by upwards of 5mm in length and 3mm in diameter. This will allow users to adapt the system to the locally available parts. Additionally the modularity of the motor system allows for the deployment of locally available motors without significant modification to the systems design.Electronics Package Makerbot Gen 4 TechZone Electronics costs 3 motors 5 motors 3 axis system 5 motor system370 103 145 473 515240 63 105 303 345JrKerr SnapMotors 160 480 800 640 960Servo Electronics 315 18 30 333 345Figure 2: Costs of 3-axis electronics plus motors from various suppliersFigure 3: Model 3 Chassis and Transmission Bill of Materials131Results: Ease of ModificationThe development of motor bays, allows the same chassis to be coupled to severaldifferent types of motor assemblies. In Figure 4, we see a Model 3 with servomotors and an identical chassis with a SnapMotors attached to the system.

Stepper motors(a) and servo motors(b) require slightly different y-axis configurations. Results: Expanded Functionalityabcdefi Figure 6: Tool heads which work on the Model 3 Fab@Home include: a Dremel mount (a), a dual-syringe tool(b), Foam cutter(c), plastic deposition(d), 4 material pressure driven tool(e), a vinyl cutter(f), a USB microscope(h), a pipetting tool(i), and a pen plotter(j)hjThe expansive variety of tool heads that are now compatible with the Fab@Home system will ensure that anyone interested in fabrication will be able to find a use for a Model 3. In figure 6 we can see nine tool heads, two of them use servos in the actuation system and would only be suitable for Model 3 system. Two of them use JrKerr Snap motors and would be compatible with Model 3 or 2.

Two of them use JrKerr Snap motors and would be compatible with Model 3 or 2. The remainders contain no onboard actuation, and are compatible with Model 2 and 3.abFigure 7: Fab@School Safety additions include a Safety hood (a) which shields the user from the machine and vents the gasses from the printer through activated carbon, and added compliance in the belt drive that limits the force the machine can apply. A spring system(b) applied a tension to the belt proportional to the compression of the spring. If the driving motor applies a force greater than the springs, the spring will compress and cause the belt to skip.The Model 3 can easily be modified to meet the safety requirement of an elementary school. The safety hood, seen in figure 7a, has been successfully fielded with students in the UVA pilot program. This system allows the Fab@Home to be the first low cost 3D printer which can be safely used in an elementary school environment.

This system allows the Fab@Home to be the first low cost 3D printer which can be safely used in an elementary school environment. The acrylic walls prevent students from being harmed by the moving parts, and the ventilation system ensures that all of the fumes pass through activated charcoal to contain the smells from non-toxic processes. The system is hinged at all joints between the vertical walls, and the top plate latches into place, making the system very portable. The added compliance from the spring based tensioning system ensure that the system cannot apply any force over a set maximum. When the limit is reached, the spring compresses and the belts will skip rather than apply further force on the moving assemblies.ConclusionsThe Model 3 Fab@Home system is the most versatile Personal Fabricator thanks to its plethora of tool heads, its optimization for modifiability, and its low barriers to entry.

The new modular motor design contributes significantly to the systems supply chain flexibility and decreased costs. The systems simplicity helped to reduce the price while lowering the technical barriers to entry and build time. Model 3 should help bring new users into the SFF user community. Educators and young children now have a machine that can meet their specific needs. It is the next step in making SFF towards ubiquitous.Acknowledgements:The authors would like to thank the NSF, Motorola Founation and MacArthur Foundation and HAYSTAC for funding this project.Bibliography 3D Systems. (n.d.). Retrieved 7 2, 2011, from Bits from Bytes: http://www.bitsfrombytes.com/Evan Malone, H. L. (2007). Fab@Home: The Personal Desktop Fabricator Kit. Rapid Prototyping Journal , Vol. 13, No. 4, pp.245-255.Lipton, J. I., Cohen, D., Heinz, M., Lobovsky, M., Parad, W., Bernstien, G., et al. (2009). Fab@Home Model 2: Towards Ubiquitous Personal Fabrication Devices.

b@Home: The Personal Desktop Fabricator KitEvan Malone, Hod Lipson Mechanical & Aerospace Engineering Cornell University, Ithaca NY USASolid freeform fabrication has the potential to revolutionize manufacturing, perhaps even to allow consumers to customize and manufacture goods cost-effectively in their own home. At present, the florescence of the technology is limited by a “chicken and egg” paradox. There is insufficient consumer demand and too narrow a range of applications for SFF systems to allow mass production to reduce their cost and complexity. At the same time, consumer applications for, demand for, and indeed awareness of SFF technology is limited by its high cost and complexity. We posit that just as the personal computer revolution was spurred by development of computer kits, getting SFF technology into the hands of hobbyists and hackers will simultaneously generate applications for, and improvements to it.

To this end, we have developed the Fab@Home personal SFF kit, and are developing a user-group website to promote exchange of ideas and improvements. The designs and software for Fab@Home will be freely distributed, and constructing a first generation kit should cost roughly $2000. The kit design and a working unit will be presented, along with our experience deploying a unit for public use.IntroductionUniversal manufacturing embodied as today’s freeform fabrication systems has – like universal computers – the potential to transform human society to a degree that few creations ever have. The ability to directly fabricate functional custom objects could transform the way we design, make, deliver and consume products. But not less importantly, rapid prototyping technology has the potential to redefine the designer.

But not less importantly, rapid prototyping technology has the potential to redefine the designer. By eliminating many of the barriers of resource and skill that currently prevent ordinary inventors from realizing their own ideas, fabbers can “democratize innovation” [4,1,2]. Ubiquitous automated manufacturing can thus open the door to a new class of independent designers, a marketplace of printable blueprints, and a new economy of custom products. Just like the Internet and MP3’s have freed musical talent from control of big labels, so can widespread RP divorce technological innovation from the control of big corporations.Despite the formidable potential of rapid prototyping technology, its acceptance over the last two decades has remained disappointingly slow [8]. At present SFF systems remain very expensive and complex, focused on production of mechanical parts, and used primarily by corporate engineers, designers, and architects for prototyping and visualization.

These factors are linked in a vicious cycle which slows the development of the technology: Niche applications imply a small demand for machines, while small demand for machines keeps the machines costly and complex, limiting them to niche applications. Alternatively, if one could provide either a large market for SFF machines and products or a simple and cheap SFF machine with which end users could invent products and applications, then this same feedback coupling could instead drive a rapid expansion in SFF technology and applications.668Learning from the history of the computer revolution In attempt to break the vicious cycle of expensive equipment and niche applications, there are many lessons to be learned from the rise and growth of an equivalently universal technology: The computer. The parallels between universal computation technology and universal manufacturing technologies are astounding.

The parallels between universal computation technology and universal manufacturing technologies are astounding. Though the universal computer in its modern architecture was realized in the 40’s [5], two decades passed before it reached any significant commercial acceptance. Early inventors themselves could not foresee its huge potential, notoriously anticipating a need “for as many as five or six machines in the US” [5]. The early commercial mainframes of the ‘60s were used mostly for niche applications such as payroll and military calculations. Like today’s rapid prototyping machines, early mainframes cost tens and hundreds of thousands of dollars, required hours to complete a single job, had the size of a large refrigerator and required trained technicians to operate and maintain.(a)(b)(c)Figure 1. Early home computers trying to break into home market.

Early home computers trying to break into home market. (a) The Honeywell Kitchen Computer cost $7000 and targeted the cooking as the ‘killer app’; (b, c) The general purpose Altair 8800, credited as starting the home computer revolution, came as a $400 kit [6]669Though it was clear to early manufacturers that the home market offered great potential, it was unclear how to successfully capture that market. Early attempts of the computer industry to break into the home market through niche ‘killer apps’ failed miserably: Some brands targeted niche domains such as Honeywell’s “kitchen Computer” geared towards recipes (Figure 1a). Its high cost and narrow application prevented it from success. Though several other home computers came out in the early ‘70s [7], the MITS Altair 8800 (Figure 1b,c) is generally credit with sparking the home computer revolution.

Though several other home computers came out in the early ‘70s [7], the MITS Altair 8800 (Figure 1b,c) is generally credit with sparking the home computer revolution. Designed and sold through Popular Electronics as a $400 kit, the Alltair 8800 broke the chicken-and-egg cycle: Hobbyists and experts could now afford to dabble with computers, develop and exchange software and numerous hardware accessory projects. Computer history had entered its exponential growth era.Based on this history, it seems reasonable to imagine a low-cost multi-material SFF system in one’s home, which could produce objects or even complete integrated devices from designs which are shared or purchased online [2]. Should such systems become as available as personal computers or printers are today, the invention and personalization of small devices could become as ubiquitous as music sharing is today.

Should such systems become as available as personal computers or printers are today, the invention and personalization of small devices could become as ubiquitous as music sharing is today. MIT’s FabLab project [1] provides ample evidence that providing people with automated fabrication tools serves as an innovation catalyst; ordinary folk, with seemingly no technical background quickly learn to exploit these tools to design and realize new inventions.The Fab@Home projectInspired by this history, and the approach employed by Bowyer toward the goal of self-replicating manufacturing systems [3], we are developing an open-source, low- cost, personal SFF system kit, which we call “Fab@Home”. The aim of this project is to put SFF technology into the hands of those same curious, inventive, and entrepreneurial citizens.

The aim of this project is to put SFF technology into the hands of those same curious, inventive, and entrepreneurial citizens. In addition, taking a page from more modern history, we are developing a “wiki” web site through which we hope to inspire users of Fab@Home to exchange their ideas for applications and their improvements to the hardware and software with us and each other. We have built three complete Fab@Home systems, and two of them have been delivered to users outside of Cornell University – one to the AMTS FabLab in Pretoria, South Africa, and one to a life-sciences laboratory in Rockefeller University, New York City. The Fab@Home design, and preliminary conclusions about these first deployments are presented below.Design Fab@Home (Figure 2) is a 3-axis Cartesian gantry positioning system driven by stepper motors attached to lead screws.

Material deposition tools are modular, and the first tool we have designed is a syringe-based extrusion tool, which uses a linear stepper motor to control the syringe plunger position. The electronics and firmware of the current version of the system provide for up to 6 axes of bipolar stepper motor control at 24V, with 2 limit switches per axis of positioning plus an optional one limit switch for the remaining 3 axes. A microcontroller controls the positioning of the axes and is in bidirectional USB communications with the PC. An application running on the PC displays the real-time state of the machine numerically and graphically, and allows the user to manually670position the axes, import and perform basic modifications to STL geometry data, apply specific material properties to each STL, and to generate and execute tool paths.Figure 2.

3D CAD model of an assembled Fab@Home machineOne concern in developing our design is that there is probably a minimum threshold of quality required in new technology kit for hobbyists, below which the excitement of the new technology will be masked by the malfunctions, maintenance problems, and poor aesthetics. Users must have a sense of what the technology is capable of before they can grasp how to modify it and apply it to their own purposes. Our first design has therefore focused more on ease of use, reliability, and aesthetics than on minimizing cost.We have tried to assume a modest availability of technical tools and skills for the envisioned end user.

The person interested in our kit would need to have a laptop or PC with a USB port, and basic assembly skills and tools, including soldering and crimping of cables and connectors, but as much as possible we have tried to design the kit such that it can be fully assembled with standard components and tools such as Allen keys, screwdrivers, scissors, pliers, and soldering iron. The main assembly tasks are snapping together the acrylic structure, inserting nuts and screws and threaded inserts, bolting together the positioning system components and mounting them to the structure, making cables to connect the microcontroller to the amplifier boards, and the motors to the amplifier boards, mounting the electronic boards to the chassis, and bundling and routing of cables. The user is expected to have some patience as well: completely assembling a kit from parts to operation requires roughly 12 hours of labor.

The user is expected to have some patience as well: completely assembling a kit from parts to operation requires roughly 12 hours of labor. Currently, the parts cost is estimated to be $2000, excluding the cost of having acrylic parts laser cut, and not including shipping costs. Table I summarizes the kit hardware parts and costs.671Table I. Part types and cost breakdownCategoryMotors Electronics Bearings/Transmission Fasteners/Hardware Acrylic Control Software TotalsTotal # of Items / # of Diff. Types 4/4 25/15 53/18 360/23 5/2 3/3 447/62Materials Cost$556.70 $488.06 $723.86 $39.63 $228.81 $0.00 $2037.06Structure The structural components of our system are built from laser-cut acrylic sheet parts held together with snap-fit joinery and simple “T-nut” style screw/nut fastening, in which a square or hex nut is inserted into a slot in one part, and a screw threads in through a hole in a perpendicular part (Figure 3).

The tight manufacturing tolerances achievable with laser cut acrylic enable us to produce good orthogonality and alignment in the machine base, increasing the ease of setup and reliability of the machine. Our first three units have been produced in-house with an Epilog Helix 35W laser engraver, cutting parts directly from SolidWorks drawing files (Figure 3), which are in turn generated from a SolidWorks 3D model of the complete system.Screw through perpendicular partSquare nut inserted in slotFigure 3. (left) Drawing of kit parts arranged for laser cutting from acrylic sheet stock; (right) “T- Nut” fastening of perpendicular acrylic sheet partsThere are 33 acrylic sheet parts in the chassis, plus an additional 7 acrylic parts for the syringe tool. Currently, six sheets of 18” X 24” X 0.236” cast acrylic are used to produce the acrylic parts. The packing of parts into these sheets is not optimal – only about 50% of the sheet surface area is converted to parts.

The packing of parts into these sheets is not optimal – only about 50% of the sheet surface area is converted to parts. More efficient packing is feasible.Cutting the parts for an entire machine requires three hours on our 35W laser cutter. We are in the process of obtaining quotes for contract manufacturing of these parts to evaluate the additional cost associated with manufacturing them. An additional benefit to having the structural parts laser cut, is that there is a large installed base of laser cutters/engravers in the sign making, and trophy and gift engraving industries. The672custom nature of the work in these industries should make them amenable to kit builder’s approaching them to have parts made. In addition, most laser cutters used in these industries do not require specialized CAM software – they operate from an application printer driver, so almost any image editing or vector drawing program can be used to design or modify designs for laser cutting and engraving.

Thus, the Fab@Home kit hacker does not need to invest in 3D CAD software to make or publish new hardware designs, and the bitmap engraving possible with these laser cutters allows them to customize the appearance of their machine with images and text. It is our hope that we can negotiate some reduced cost for laser cutting services for the Fab@Home community, at least for a standard design, in exchange for advertising on the Fab@Home website and a large volume of customers.PositioningThe linear motion components of the positioning system use off-the-shelf linear ball-bearing pillow blocks running on ½” diameter rails (McMaster-Carr, Inc.). The X and Y axes are in a gantry configuration with the deposition tool riding on the Y axis, which in turn rides on the X. The Z axis moves the build surface independently from the X and Y to minimize acceleration of parts as they are being fabricated. We have selected HSI Inc.

The Z axis moves the build surface independently from the X and Y to minimize acceleration of parts as they are being fabricated. We have selected HSI Inc. linear stepper motors for our actuators because of their simple design, high resolution, and the semi-custom manufacturing focus of the company which permits specifying precisely the leadscrew and bearing journal dimensions required for our application, simplifying the overall design and assembly. For the X,Y, and Z axes we use NEMA size 14 bipolar motors with rotor-mounted lead screws. External polymer lead nuts are mounted to the axis carriages. In the case of the X axis, a timing belt and pulleys (Stock Drive Products, Inc.) are used to couple a slave leadscrew to the motor leadscrew to achieve symmetrical drive of the gantry. The force, maximum speed, and positioning resolution all depend upon the lead screw threading selected.

The force, maximum speed, and positioning resolution all depend upon the lead screw threading selected. We use the “B” series threading for the positioning axes, which allows for 15.8 μm travel per full step, a nominal top speed of 25 mm/s (1600 step/s), and a maximum thrust of 120 N.Syringe ToolThe syringe tool structure (Figure 4) is also constructed of laser cut acrylic parts with snap fit joinery and T-nut fasteners. A linear stepper motor controls the position of the syringe piston. We employ as NEMA size 8 frame motor with a rotor mounted lead nut. The lead screw, which is not captive in the motor, has “7” series threading, 3.2 μm travel per full step, and the motor can achieve a top speed of 5.8 mm/s (1800 step/s), and a maximum thrust of 90 N. For the 10cc syringes we use, this amounts to a 1.1 cc/s maximum volume flow rate, and a maximum syringe pressure of 460 kPa (67 PSI). The current syringe tool has been designed to allow EFD Inc.

For the 10cc syringes we use, this amounts to a 1.1 cc/s maximum volume flow rate, and a maximum syringe pressure of 460 kPa (67 PSI). The current syringe tool has been designed to allow EFD Inc. 10cc disposable syringe barrels to snap in and out, and for the piston to be quickly attached and released from the motor leadscrew for quick changing of materials. A metal nut fits tightly inside of disposable syringe pistons, and one end of the motor leadscrew has threading to match the nut. When firmly threaded into the nut, the leadscrew is prevented from rotating with the motor rotor, and hence the rotor motion is converted to linear motion. Manually unscrewing the leadscrew from the nut allows exchanging syringes regardless of how full673without the need to move or remove the piston, facilitating fabrication of multiple- material objects, and conserving materials.Figure 4.

(left) The standard design, single syringe tool, driven by a linear stepper motor; (right) A two-syringe version for a life-sciences laboratoryWe have also developed a dual syringe tool (Figure 4) which allows two materials to be loaded simultaneously and independently deposited. As mentioned before, tools are bolted to the positioning system, and tools are intended to be modular. We have selected a syringe deposition tool as the first design because of the broad range of materials useable with such tools, and for the intuitiveness of operation.ElectronicsPersonal computers today typically provide several USB connections, but no longer have RS-232 serial ports, though this is still the de facto standard serial communication method for robotics and microcontroller hobbyists. As a result, we have opted to support direct USB connection to our Fab@Home system, despite the additional development work and (internal) complexity that this entails.

As a result, we have opted to support direct USB connection to our Fab@Home system, despite the additional development work and (internal) complexity that this entails. We chose to use a microcontroller with an on-chip USB 2.0 peripheral, the Philips LPC-2148 ARM7TDMI (Royal Philips Electronics N.V.). This is a very high performance, 60MHz, flash- memory microcontroller with a wealth of peripheral functions for future expansion, including ADC, DAC, PWM, counter/timers, real-time clock, high-speed GPIO, UARTs, SPI, I2C, not to mention the USB2.0 peripheral. In addition, it has 512kB of flash memory, and 40kB of RAM. The large program memory has enabled us to make a very easily understood and extended packet data protocol for communication between the PC application and the firmware. We use the large RAM space to buffer motion commands so that real-time motion does not depend on variations in communication bandwidth.

We use the large RAM space to buffer motion commands so that real-time motion does not depend on variations in communication bandwidth. With our current protocol, we can buffer roughly 670 path points for 6 axes of control. The microcontroller is powered by the USB, and thus can be communicated with even674when the amplifier electronics are not powered. The high computational performance of the device enables the system to handle receiving and buffering path points, sending real- time status and position data, and controlling step and direction outputs for 6 axes at least 5kHz – we have not yet tested higher step rates. The microcontroller is available on a 1.5” X 2.5” board (Figure 5) with header connectors for all pins and a USB connector for $39.95 in single quantity (LPC-H2148, Olimex Inc.).Figure 5.

The microcontroller is available on a 1.5” X 2.5” board (Figure 5) with header connectors for all pins and a USB connector for $39.95 in single quantity (LPC-H2148, Olimex Inc.).Figure 5. The electronics boards – two 3-axis stepper motor amplifiers on the left, LPC-H2148 microcontroller boards on rightCurrently, we are using two 3-axis stepper motor amplifier boards (Figure 5) ($170 ea., Technological Arts, Inc.) to power the positioning system and syringe tool stepper motors. These boards have proven to be problematic, and we have had to upgrade some of the components on them to make them function reliably. These boards provide switch-mode current regulation for up to 3 bipolar stepper motors per board. The current regulation allows us to use 24VDC supply and 5V rated motors to get much higher acceleration (hence faster builds at finer resolutions), than would be possible at the nominal 5V.

The current regulation allows us to use 24VDC supply and 5V rated motors to get much higher acceleration (hence faster builds at finer resolutions), than would be possible at the nominal 5V. This is a design decision which increases cost significantly for the sake of making the technology more useable. We have found and are evaluating alternate amplifier boards which provide the same or improved functionality at less than half the price.Software The firmware for the LPC-2148 microcontroller was developed in C language, using the free GNU C ARM compiler, and Rowley CrossWorks for ARM development environment by Rowley Co. UK – available for 30 days unrestricted trial, and with a 99£ educational license, and a 499£ commercial license. CrossWorks is not essential - several freeware IDE’s exist which work with GNU C ARM.

The firmware performs the following main functions:(cid:120)receiving and parsing of packetized commands from the PC via the USB675(cid:120) (cid:120) (cid:120) (cid:120)(cid:120)buffering of motion path segments for fabrication paths immediate execution of jog motion commands configuration of limit switches (present/absent for each axis and direction) communicating axis position, limit switch state, and other system status to the PC via the USB controlling step and direction outputs for up to 6 axes at > 5kHz step frequencyAs mentioned before, the microcontroller has additional resources available for future expansion, and the firmware has been designed with ease of expansion in mind.A PC application has been written which enables the user to control the machine, import, position, assign material properties to, and generate and execute manufacturing plans for geometry data which is imported in the form of STL files.

In addition, the application has been designed around the concept that while the core application may eventually be improved by the user community, in the short term, the hardware configuration, the types of materials, and parameters for depositing materials would be far simpler and more interesting to explore and share. Thus the hardware configuration, material properties, and material deposition parameters are described in plain text parameter files. These files describe, for instance, the color and geometry data used to render the Fab@Home machine, the speeds and threading of the stepper motors, presence of limit switches, etc.The PC application is targeted for the Microsoft Windows operating system. It is written in C++ using the Microsoft Visual Studio .NET development environment, Open GL for graphics rendering, and the Microsoft Foundation Class library for user interface components.

It is written in C++ using the Microsoft Visual Studio .NET development environment, Open GL for graphics rendering, and the Microsoft Foundation Class library for user interface components. The application has been designed with the aim of maximizing the intuitiveness of use (Figure 6). The user interface includes a 3D rendering of a Fab@Home machine which moves synchronously with the real-time position information sent back by the microcontroller. Dialog boxes allow importing and assigning material and tool properties to the part geometry, manual jogging of the axes via buttons and mouse scroll wheel, including the syringe tool motor, as well as a numerical view of the real-time position and status data from the microcontroller. The application also allows a rudimentary simulation of the fabrication process – the actual manufacturing plan is executed on a Fab@Home software emulator, and the motions are displayed in the GUI for quick checking of toolpaths.676Figure 6.

A screenshot from the PC application displaying a model ready for fabrication and dialog boxes for positioning and real-time status informationThe workflow for Fab@Home consists of the following:(cid:120)(cid:120)(cid:120) (cid:120) (cid:120)(cid:120) (cid:120) (cid:120) (cid:120) (cid:120) (cid:120) (cid:120) (cid:120)connecting PC to Fab@Home via USB cable, and plugging in the Fab@Home power supply selecting and/or modifying/tuning parameter files to match the hardware and materials to be used starting the PC application loading the parameter files loading a syringe with piston, material, and nozzle, and mounting it in the tool; threading the tool leadscrew into the piston nut importing the geometry of the part to be fabricated assigning material and tool properties to part geometry automatic generation of a manufacturing plan if desired, simulated execution of the plan initialization of the Fab@Home printer homing of the axes so that GUI and physical positions match jogging axes to the desired origin for fabrication automatic execution of the manufacturing planResultsTo date, we have produced three complete Fab@Home units, the first two ofwhich have been delivered to users outside of Cornell University.The very first unit was delivered to the AMTS FabLab in Pretoria, South Africa, on June 25, 2006.

The AMTS FabLab is a publicly accessible CAD/CAM lab developed by the MIT Center for Bits and Atoms in partnership with the South African Government’s Department of Science and Technology’s Advanced Manufacturing Technology Strategy and the South African Council for Scientific and Industrial Research (CSIR). FabLab’s, the invention of Prof. Neil Gershenfeld of MIT, are intended677to provide simple, highly automated CAD/CAM facilities and basic training to members of the general public in order to promote local invention and micro-enterprise – “Democratization of Invention”. During the Digital Manufacturing Symposium held at CSIR during the week of June 25th, we were able to demonstrate Fab@Home to officials of the CSIR and AMTS as well as to visitors and staff of the FabLab. Unfortunately, it being the very first fully functional unit, both the software and the hardware had significant bugs, and demonstrations were limited to the deposition of single layers.

Unfortunately, it being the very first fully functional unit, both the software and the hardware had significant bugs, and demonstrations were limited to the deposition of single layers. Nevertheless, viewers were thrilled by the idea of a low-cost SFF system, and also with the visual appearance of the software and hardware. Gershenfeld (pers. comm.) views the Fab@Home kit as an example of a class of CAM equipment which can be cheaply manufactured using a the simple CAM tools available in a FabLab, with the result that FabLabs may be able to extend their own capabilities, and expand to have greater impact with little additional capital expense. The AMTS has committed resources to integrating Fab@Home into the CAM suite at the FabLab, and is using Fab@Home as a prototype for generating additional examples of kit CAM equipment.Figure 7.

The AMTS has committed resources to integrating Fab@Home into the CAM suite at the FabLab, and is using Fab@Home as a prototype for generating additional examples of kit CAM equipment.Figure 7. The first unit being demonstrated for staff and users of the Soshanguve FabLab public access CAD/CAM laboratory, Soshanguve Township, South AfricaWhile in South Africa, we also demonstrated the Fab@Home at the Soshanguve FabLab, situated in Soshanguve Township, an entirely black African slum of 750,000 residents. Here again, the viewers, adults and children, most poorly educated, were quite excited at the appearance of the machine but typically failed to differentiate it from a CNC mill – a device with which they had some experience in the FabLab. A noteworthy reaction was that Fab@Home was an example of a cheap way to build a CNC mill in the FabLab.

A noteworthy reaction was that Fab@Home was an example of a cheap way to build a CNC mill in the FabLab. At this demonstration, we may have witnessed the aforementioned threshold of quality – the impact of the technology was lost for the lack of smooth operation and a meaningful example application. Nonetheless, the managers of the lab expressed and urgent desire for a Fab@Home unit for their lab – a request which we will fill in the near future.678We have very recently produced and delivered a Fab@Home unit (Figure 8) to a life science research laboratory at Rockefeller University, New York, NY. As of this unit, the hardware and firmware have been satisfactorily debugged. At this laboratory, research is underway which involves the slime mold organism Dictostylium, which transitions from single-cell independent life to colony organism with coordinated motion in response to scarcity of nutrition.

The Fab@Home offers a uniquely simple and customizable platform for executing experiments on the effect of initial spatial distribution of Dictostylium cells in the environment and the spatiotemporal evolution of their colony aggregation response to starvation stress. A dual syringe tool was designed and included with this unit to facilitate the spatial control of both organisms and nutrients in experimental geometries. We are eagerly awaiting preliminary results from this first scientific application of Fab@Home.Figure 8. The second unit with a 2-syringe tool deployed by research student for Dictostylium research at Rockefeller University, NYC, NYA third unit (Figure 9) has been assembled and is being used for ongoing debugging and development of the software.

Minor hardware enhancements have been added to simplify the routing and strain-relieving of cables, and the hardware assembly process is becoming more routine.The software has been improved to the point that bugs are no longer the most obvious limiting factor in performance. As Figure 9 shows, thin-walled, and multilayer parts are now achievable with RTV silicone rubber, but the part quality can be improved by more careful tuning of the deposition and machine control parameters.679Figure 9. (left) The 3rd unit undergoing testing at Cornell University; (right) A successful part built by a Fab@Home using 1-part RTV silicone rubberFuture WorkAlthough we have reserved internet URL’s for our website, we have only just begun the development of our user group website. The RepRap website [3] provides a nice example of the type of functionality we intend to implement in our website, including the user editability provided by a “wiki” framework.

The RepRap website [3] provides a nice example of the type of functionality we intend to implement in our website, including the user editability provided by a “wiki” framework. We are awaiting a first round of bids for outsourced manufacturing (MFG.COM) of the laser cut acrylic parts, which we will use to improve quality of our cost estimate for the kit. While writing software documentation and continuing improvements in the software, we will recruit volunteers at Cornell University to assemble units from parts in order to simplify, debug, and document the assembly procedure, parts lists, and source lists for the hardware. We also intend to have these volunteers and our first two users begin to populate the website with applications, tips, modifications, and bug reports in the interest of jump-starting a user community. Also in this interest, we are considering giving away a few more units to selected candidate users, such as the Soshanguve FabLab.

Also in this interest, we are considering giving away a few more units to selected candidate users, such as the Soshanguve FabLab. In the longer term, we will begin to optimize the design to minimize cost, and enhance performance, perhaps developing design variants.Conclusions Few technologies can truly transform human society to a degree that computers have. The universal computer, from the early experimental machines of the ‘50s to the home computer today, has affected almost every aspect of human existence. It is not easy to foresee such disruptive technologies in advance, but it has not escaped many that solid freeform fabrication technologies bare many of the same traits.

It is not easy to foresee such disruptive technologies in advance, but it has not escaped many that solid freeform fabrication technologies bare many of the same traits. Like universal computation, universal manufacturing, embodied as today’s freeform fabrication systems, has the potential to transform human society to a degree that few creations ever have.In order to accelerate the spread and development of SFF applications and technology, and escape the “chicken and egg” paradox, we are developing an open- source, low-cost, personal SFF system kit, which we call “Fab@Home”. The current kit680design has a parts cost of roughly $2000, requires only basic hobbyist tools and skills to assembly and use, and can be used to deposit almost any room temperature liquid or paste. We are also developing a website to foster a user community and to promote the exchange of ideas and improvements of the technology.

We are also developing a website to foster a user community and to promote the exchange of ideas and improvements of the technology. We have thus far built three complete Fab@Home systems, and two of them have been delivered to users outside of Cornell University. The response of users, government officials, and the general public has been very positive, despite the problems associated with deploying prototype systems. Generating an initial wave of interest does not seem to be a concern – what remains to be seen is whether we can provide a system cheap enough and simple enough that for many people, the experience of obtaining and operating an SFF system does not overwhelm the experience of using and exploring the technology.Bibliography 1. Gershenfeld N., (2005) FAB: The Coming Revolution on Your Desktop – FromPersonal Computers to Personal Fabrication, Basic Books2. Lipson H.

Presented in this paper is a further development of this process focusing on fabrication of functionally graded materials (FGM). A dynamic mixing mechanism was developed for mixing constituent ceramic pastes, and an extrusion control scheme was developed for fabricating specimens with desired material compositions graded in real time. FGM specimens with compositions graded between Al2O3 and ZrO2 were fabricated and ultimately densified by sintering to validate the effectiveness of the CODE process for FGM fabrication. Energy dispersive spectroscopy (EDS) was used to compare final compositions to the original material designs. The specimen’s hardness at different locations along the gradients was examined by micro-indentation tests.

The specimen’s hardness at different locations along the gradients was examined by micro-indentation tests. The dimensions of sintered specimens were measured, and the effects of material composition gradients on the distortions of sintered FGM specimens were analyzed.IntroductionSeveral additive manufacturing (AM) processes have been developed for ceramics and glasses, including binder jetting [1], [2], material extrusion [3]–[8], vat photopolymerization [9], [10], powder bed fusion [11, 12], and directed energy deposition [13–15], among others. Ceramic On-Demand Extrusion (CODE) [7] is a recently developed paste extrusion-based AM process, which produces ceramic components with near theoretical density (>98%) after sintering [4, 16– 18]. It deposits high solids loading (>50 vol%) aqueous ceramic pastes onto a substrate layer by layer at room temperature.

It deposits high solids loading (>50 vol%) aqueous ceramic pastes onto a substrate layer by layer at room temperature. Each deposited layer is partially solidified by uniform infrared radiation drying from above before the subsequent layer is initiated. At the same time, undesirable evaporation from the sides of the part is prohibited by surrounding the part with a liquid [7]. This layered uniform radiation drying approach minimizes the moisture content gradient in the part1087during the fabrication process and thus enables CODE to produce crack-free ceramic parts.

This layered uniform radiation drying approach minimizes the moisture content gradient in the part1087during the fabrication process and thus enables CODE to produce crack-free ceramic parts. The printed parts are then bulk-dried in a controlled environment with appropriate humidity, after which the green bodies are sintered to produce near-theoretical density parts [16–18].Functionally graded materials (FGM) are characterized by gradual variations of material compositions over volumes, which allows for a combination of materials or material properties not typically achievable in monolithic materials [19]–[22]. Alumina/zirconia (Al2O3/ZrO2) FGM components have been of great interest, mainly to enhance the fracture toughness through the incorporation of ZrO2 phase.

Alumina/zirconia (Al2O3/ZrO2) FGM components have been of great interest, mainly to enhance the fracture toughness through the incorporation of ZrO2 phase. One potential application is the prosthetic hip joint ball [22, 23], where the FGM realizes the transition from a tough ZrO2 core, which provides the high strength of the hub structure and reduces the risk of cracking, to a wear-resistant Al2O3 ball surface, which guarantees a long service life in a human body. Additive manufacturing (AM) processes are especially advantageous for fabricating FGM components due to the layer-by-layer nature of the processes.

Additive manufacturing (AM) processes are especially advantageous for fabricating FGM components due to the layer-by-layer nature of the processes. Considering that the melting temperatures of ceramics are usually too high for thermal- based melt deposition and the fact that the ink jetting-based ceramic AM processes are subject to high porosity [20], material extrusion-based AM processes are the most favorable method for fabricating ceramic FGM components [20].The present paper introduces a dynamic mixing device to the Ceramic On-Demand Extrusion (CODE) system for fabricating FGM specimens. Two distinct materials were extruded through separate extruders into the mixing chamber of the dynamic mixer with controlled flowrates. The pastes were then blended by the dynamic mixer to a homogenous mixture and finally deposited through nozzles to fabricate FGM specimens with planned material compositional distribution.

The pastes were then blended by the dynamic mixer to a homogenous mixture and finally deposited through nozzles to fabricate FGM specimens with planned material compositional distribution. The FGM specimens were post-processed and characterized to validate the functionality and accuracy of the dynamic mixing device. Deformation was observed on the sintered specimens. The effects of material composition gradients on the deformation of sintered FGM specimens were analyzed.Experimental SetupThe CODE fabrication system consists of a cartesian gantry system, an extrusion device mounted on the gantry and capable of extruding viscous ceramic pastes at controlled flowrates, an oil feeding device to regulate the oil level in the tank, and an infrared radiation heating device capable of positioning the infrared source and providing on/off control. Introduction and details of the CODE fabrication system were presented in the previous papers [4, 7, 24].

Introduction and details of the CODE fabrication system were presented in the previous papers [4, 7, 24]. In the present work, the CODE system was configured to deposit a mixture of two materials using a dynamic mixing tool head consisting of two auger extruders and a dynamic mixer. The two auger extruders were controlled separately to extrude two different pastes at independent flowrates. As shown in Figure 1(a), the two pastes are extruded through the inlets into the dynamic mixer. As the two pastes passed through the mixing chamber, a motor-driven mixing blade blended the distinct pastes to produce a homogenous mixture, which was then deposited through the outlet to form a 3D part. The actual components and layout of the dynamic mixing tool head are shown in Figure 1(b).1088Figure 1. Dynamic mixing tool head of the CODE system.

The actual components and layout of the dynamic mixing tool head are shown in Figure 1(b).1088Figure 1. Dynamic mixing tool head of the CODE system. (a) Schematic of the dynamic mixer, (b) The actual dynamic mixing tool head mounted on the CODE system.Fabrication of Functionally Graded MaterialsPreparation of pastesAqueous Al2O3 pastes and ZrO2 (3Y) pastes were prepared as the feedstock materials. The composition of the pastes and detailed steps of paste formation can be found in previous works [16, 17]. The ZrO2 pastes were prepared to have 50 vol% solids loading. Considering the shrinkage of components during the bulk drying process and the sintering process, the Al2O3 pastes were also adjusted to have 50 vol% solids loading, expecting to minimize the mismatch of shrinkage between the two materials, and hence minimize the distortion of the FGM specimens during sintering.

In order to visualize the material variation in the printed specimens, the white Al2O3 pastes and ZrO2 pastes were colored green and pink, respectively, using FD&C Yellow 5 & Blue 1 and FD&C Red 40 dyes.Printing single-bead lines with graded materialsIn order to test the capability of the dynamic mixing tool head to build parts with graded materials and examine the response of changing material composition, a single-bead serpentine shape was printed using both the ZrO2 paste (pink) and the Al2O3 paste (green). The ZrO2 and Al2O3 pastes were extruded by extruders A and B, respectively. As shown in Figure 2, the dynamic mixing tool head was first loaded with ZrO2 paste and was then commanded to switch to Al2O3 paste by turning off extruder A and turning on the extruder B at location 1. The serpentine shape was printed from location 1 to location 5, during which the dynamic mixing tool head was commanded to switch between ZrO2 and Al2O3 at locations 2, 3, and 4.

The serpentine shape was printed from location 1 to location 5, during which the dynamic mixing tool head was commanded to switch between ZrO2 and Al2O3 at locations 2, 3, and 4. The mixing blade was spinning at 900 rpm during the entire printing process. As expected, the change in material didn’t1089take place immediately after receiving the control commands at locations 1-4. Instead, a delay was observed before the printed line changed its color.Figure 2. A single-bead serpentine printed using dynamic mixing device with graded materials (pink: ZrO2,green: Al2O3). The mixing blade was spinning at 900 rpm during the entire printing process.Determining the delay of changing material compositionThe dried serpentine shaped specimen is shown in Figure 2 and Figure 3(a). The colors in Figure 3(a) indicated that the material composition changed periodically in four cycles as commanded.

The colors in Figure 3(a) indicated that the material composition changed periodically in four cycles as commanded. The pattern of material color change was similar in each cycle, indicating that the delay in composition transition was repeatable. Other than observing the material composition variation visually by color, EDS was also utilized on sintered specimens to obtain detailed measurements of composition at locations of interest. As shown in Figure 3(b), 42 locations of interest in cycle 1 were selected for EDS measurements. Locations 1 and 42 were the starting and ending locations of cycle 1, i.e., the locations where the tool head was commanded to switch from one material to another. The EDS measurements were taken at the surface of the specimen. At each location, the EDS measurement covered three areas of ~ 120 µm by 120 µm.

The EDS measurements were taken at the surface of the specimen. At each location, the EDS measurement covered three areas of ~ 120 µm by 120 µm. The percentages of Al atomic counts over the total atomic counts of Al, Zr, and Y of the three sampling areas for all locations are plotted in Figure 3(c). According to the graph in Figure 3(c), after the command to change material was received at location 1, the actual material composition started to change at location 7, and gradually transit to the desired material composition (100% Al2O3). Location 38 was determined as the location where the transition of composition was completed. The stable material composition between locations 1 and 7 was characterized as the result of purging out the residual ZrO2 paste in the material outlet (shown in Figure 1(a)). The volume of paste being deposited between locations 1 and 7 was termed the delivery delay.

The volume of paste being deposited between locations 1 and 7 was termed the delivery delay. On the other hand, the volume of paste being deposited between locations 7 and 38 was termed the transition delay, which was related to the paste streams propagating behavior in the flow path of the mixer, and the volume of the flow path. The delivery delay and the transition delay made up the total delay when changing material composition, and their volumes were quantified as 0.12 mL, 0.36 mL, and 0.48 mL,1090respectively. As expected, the delivery delay (0.12 mL) was equal to the volume of the material outlet of the mixer. In the practice of printing laminar FGM samples, the total delay (Vdelay) was considered in the control scheme, and proper compensation was made to precisely deposit materials with desired compositions.Figure 3. (a) Four cycles of the printed serpentine. (b) 42 sampling locations for EDS test on cycle 1 of the serpentine.

(a) Four cycles of the printed serpentine. (b) 42 sampling locations for EDS test on cycle 1 of the serpentine. (c) The atomic percentage of Al measured by EDS at different locations.Fabricating laminar Al2O3/ ZrO2 FGM specimensTo examine the ability of the dynamic mixing tool head to accurately control the material composition, and to study the drying and sintering behavior of bulk FGM specimens, three groups of laminar FGM specimens graded from Al2O3 to Al2O3/ZrO2 (50/50) were fabricated. Each group had three samples with different material gradients. The material gradient for the three types is shown in Figure 4. Type 1 had a 5 vol% increment of ZrO2 every one layer and hence grads from Al2O3 to Al2O3/ZrO2 by 10 steps through 11 layers. Type 2 had a 10 vol% increment of ZrO2 every 2 layers hence grads from Al2O3 to Al2O3/ZrO2 by 5 steps through 11 layers. Finally, Type 3 was transitioned from Al2O3 directly to Al2O3/ZrO2 at the 6th layer of the 11-layer sample.

Finally, Type 3 was transitioned from Al2O3 directly to Al2O3/ZrO2 at the 6th layer of the 11-layer sample. All the FGM specimens were designed to have identical dimensions (70×20×6.6 mm3) and numbers of layers to ensure a fair comparison of potential part deformation after sintering.1091Figure 4. Design of material composition distribution in three types of FGM specimens.Considering the significant delay (Vdelay = 0.48 mL) of changing composition discussed in the previous section, each time a change of composition in a new layer was required, an additional purging process was conducted before the new layer was started. In the purging process, the dynamic mixing tool head was commanded to move out of the building substrate, extrude two pastes at predetermined flowrates for a certain volume Vpurge until the materials at the nozzle tip reached the desired steady composition (Vpurge ≥ Vdelay), and was finally moved back to the building substrate to start the new layer.

The entire printing and purging process was fully automated by G- Codes. A Vpurge of 0.5 mL was used in this experiment. The printed FGM specimens were bulk dried in an environmental chamber for 12 hours where the temperature was controlled at 24°C and the relative humidity at 75%, then fired at 500°C for 1 hour to burn out the organic constituents including binder and dyes, and finally sintered at 1500°C for 1 hour to densify. The sintered specimens were characterized, and the result is presented in the following section. Figure 5 shows the printing process of an FGM specimen and dried FGM specimens of three types.Figure 5. (a) A laminar Al2O3/ZrO2 FGM specimen being printed using the dynamic mixing tool head.

Figure 5 shows the printing process of an FGM specimen and dried FGM specimens of three types.Figure 5. (a) A laminar Al2O3/ZrO2 FGM specimen being printed using the dynamic mixing tool head. (b)(c) (d) A group of laminar Al2O3/ZrO2 FGM specimens printed and dried.1092Results and DiscussionAccuracy of material composition controlTwo sintered FGM specimens from Type 1 were cut, and ground from the side for ~ 4 mm to reveal the inner cross-section as illustrated in Figure 6(b), then polished down to a 0.25-micron finish and coated with gold/palladium for EDS measurements. For all 11 layers in the two specimens, EDS intensity measurements were taken over 3 regions with aproximately120 µm by 120 µm areas. The ratio of Al peak vs. Zr peak, as well as the SEM images of two example layers, are shown in Figure 6(c)(d). The dark (Al2O3) and bright (ZrO2) phases in the SEM images show a homogeneous distribution of the two materials.Figure 6.

The dark (Al2O3) and bright (ZrO2) phases in the SEM images show a homogeneous distribution of the two materials.Figure 6. (a) A photo of a dried (not sintered) FGM specimen of Type 1, and a close-up view showing a color change of its layers. (b) Schematic showing the cross-section where EDS measurements were taken. (c) (d) EDS peaks and SEM images taken from layer 11 and 6 of the specimen.The mean values of the measured atomic percentage of Al for layers 1-11 were plotted in Figure 7 by the solid line. Based on the designed volume ratio of the Al2O3 and ZrO2 pastes, the solids loading of two pastes, density and molecular weight of the two materials, the nominal atomic percentages of Al in all layers were calculated and plotted in Figure 7 by the dashed line. The average error between the measured value and nominal value was 1% and the maximum discrepancy was 2%, indicating a good accuracy of the dynamic mixing tool head for controlling material composition.1093Figure 7.

The atomic percentage of Al of each layer in the sintered FGM specimens.Deformation of specimensAfter sintering, the 3 groups of colored Al2O3/ZrO2 specimens turned white since the organic dyes burned out during the firing process. Curling and delamination were observed in the sintered specimens. In order to quantify the deformation of each specimen, as shown in Figure 8, the specimens were flipped on a flat substrate to measure the heights at their center and two ends, from which the curling angles were calculated for every specimen. The quantified deformation of all FGM specimens is reported in Table 1.Figure 8. One group of FGM specimens after sintering, deformation, and failure were observed.1094Table 1.

The quantified deformation of all FGM specimens is reported in Table 1.Figure 8. One group of FGM specimens after sintering, deformation, and failure were observed.1094Table 1. Quantified deformation (curling angle) of each FGM specimen after sintering.Grading StepGroup 1Curling AngleGroup 2Group 35% (Type 1)2.7°2.3°2.9°10% (Type 2)50% (Type 3)3.6°4.9° (cracked)3.8° delaminated3.7° delaminatedVickers hardnessA sintered FGM specimen from Type 2 was cut, ground and polished in the similar way illustrated in Figure 6(b). Micro-indentation tests were performed at layers with different material compositions following ASTM C1327 standard [26]. For each layer with different compositions, 5 indentations were performed at different locations and the mean values of Vickers hardness were presented in Table 2, as well as their standard deviations.Table 2.

Vickers hardness of layers with different composition of Al2O3/ZrO2 in an FGM specimen.LayerVolume ratio of Al2O3 vs. ZrO2Vickers Hardness (GPa)MeanStandard Deviation10-11 8-9 6-7 4-5 2-3 150%:50% 60%:40% 70%:30% 80%:20% 90%:10% 100%:0%15.0 16.0 16.3 16.9 17.5 18.40.31 0.23 0.14 0.41 0.28 0.32DiscussionDue to the delay of changing material composition using the dynamic mixing tool head, time and material waste are introduced to the printing of laminar FGM specimens when pastes are purged out before they reach a desired stable composition for a new layer. Moreover, the transition delay also limits the accuracy of controlling the in-plane gradient of composition in one layer. Minimizing the delay is always desirable for fabricating FGM components. The key point of minimizing the delay is to reduce the volume of the flow path.

Minimizing the delay is always desirable for fabricating FGM components. The key point of minimizing the delay is to reduce the volume of the flow path. However, tighter flow path leads to an increase of back pressure, which requires higher extrusion pressure, especially for highly- viscous ceramic pastes. In addition, the fabrication of micro-sized mixing blades and flow paths is also challenging. A trade-off between the fast transition of composition and equipment investment is necessary, which will be investigated in the future, as well as the overall design of the dynamic mixer.According to Figure 8 and Table 1, as the step of changing composition increased, the amount of deformation of the sintered FGM specimen increased until structural failure (cracking1095and delamination) took place. Although less significant, a similar trend of deformation was also observed on the FGM specimens after they were bulk dried, as shown in Figure 5(b)(c)(d).

Although less significant, a similar trend of deformation was also observed on the FGM specimens after they were bulk dried, as shown in Figure 5(b)(c)(d). The deformation occurred during the bulk drying process indicated a mismatch in drying shrinkage of the two pastes. Since the two pastes had the same (50 vol%) solids loading, the difference in shrinkage was believed to be caused by the different packing density of the particles in the two pastes after they were dried. On the other hand, the deformation occurred during the sintering process was believed to be caused by the mismatch of sintering shrinkage and thermal expansion of Al2O3 and ZrO2. Larger differences in material composition between layers lead to larger stresses caused by the mismatch of material properties, which explains the fact that the larger step of changing composition led to larger amounts of deformation.

A smoother (reduced) gradient of material composition is likely to reduce the amount of deformation and the risk of part failure. Adjusting the inherent properties of the raw materials to reduce the mismatch of shrinkage could be another effective way of mitigating the stress and deformation. For example, according to the study of Sun et.al. [27], adjusting the particle size distribution of the Al2O3 and ZrO2 powders could contribute to matching the shrinkage of layers with different Al2O3/ZrO2 ratios in the FGM specimens.SEM images taken at the cross-section of the FGM specimens, of which two example images are shown in Figure 6, demonstrated a high homogeneity of the mixed Al2O3 and ZrO2 phases. Vickers hardness reported in Table 2 demonstrated a clear trend of decreased hardness as the ZrO2 concentration was increased. The hardness values were in good agreement with data reported from other literature [27–29].

The hardness values were in good agreement with data reported from other literature [27–29]. The hardness measurements at different locations for each composition were highly consistent according to their standard deviations, which is another evidence that the Al2O3 and ZrO2 pastes were mixed to highly homogeneous.ConclusionsA dynamic extrusion-mixing device was developed for fabricating functionally graded materials (FGM) by the Ceramic On-Demand Extrusion process. The effectiveness of this device was validated by fabricating FGM specimens graded from pure Al2O3 to Al2O3/ZrO2 with predetermined material gradients. The FGM specimens were sintered at 1500 °C and the material composition of each layer in the sintered FGM specimens was validated by measuring the atomic percentage of Al and Zr by energy dispersive spectroscopy (EDS), and compared to the original design of compositions. A 1% average error of material composition was observed from the comparison.

A 1% average error of material composition was observed from the comparison. Deformations were measured on the dried and sintered FGM specimens. Greater material composition gradients led to larger deformations of specimens. The Vickers hardness decreased from 18.4 GPa to 15.0 GPa as the volume percentage of ZrO2 increased from 0 to 50 vol% in Al2O3/ZrO2 layers.AcknowledgmentsThis work was funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy. The United States Government1096retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid up, irrevocable, worldwide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.References[1]S. Zhang, H. Miyanaji, L. Yang, A. Ali Zandinejad, J. Dilip, and B.

bricating Superhydrophobic Surfaces with Solid Freeform Fabrication ToolsAlan M. Lyons1,2*, John Mullins2, Mark Barahman,1 Itay Erlich,1 and Todd Salamon31CUNY – College of Staten Island New York, NY, U.S.A.2Bell Laboratories – Alcatel-Lucent Dublin, Ireland3Bell Laboratories – Alcatel-Lucent Murray Hill, NJ U.S.A.Abstract: Superhydrophobic surfaces exhibit a range of properties such as large contact angle, low contact angle hysteresis and decreased hydrodynamic drag. These properties make superhydrophobic surfaces of fundamental and commercial interest as they can enable a wide variety of applications including microfluidic components, biomedical devices, and micro-batteries. Superhydrophobic behavior is achieved through a combination of the hydrophobicity of the polymer and the roughness of the surface. We have used a commercially available multi-jet modeling rapid prototyping machine to fabricate 3D objects where the superhydrophobic surface is monolithic with the part.

We have used a commercially available multi-jet modeling rapid prototyping machine to fabricate 3D objects where the superhydrophobic surface is monolithic with the part. This approach was used to fabricate non-planar components with novel structures including helical conduits and porous meshes. In addition, we have developed a robotic dispensing tool that enables greater freedom of material selection. Both approaches have been used to fabricate arrays of surface features with diameters below 175 microns and with aspect ratios greater than 8:1. The fabrication and wetting properties of surfaces made using these two techniques will be discussed.Introduction:Superhydrophobic surfaces are highly water repellant such that water droplets form acontact angle with the surface in excess of 150 degrees. This effect is due to a combination of the low solid surface free energy and extreme surface topography of the material.

This effect is due to a combination of the low solid surface free energy and extreme surface topography of the material. Drops of aqueous fluids are suspended on the tips of the structured surface elements and exhibit significant slip along the boundary. The fluid is primarily in contact with air; only a relatively small fraction (usually <5%) of the total fluid surface area is in contact with the solid surface. Drag reduction values of more than 50%, relative to a smooth surface, have been reported for flow in a water tunnel1. Reduction in flow resistance (drag) has been primarily observed for water. However, aqueous alcohol solutions have exhibited so-called superlyophobic behavior for static droplets2.Such surfaces occur naturally in both plants3 as well as animals4 and the fabrication and properties of synthetic superhydrophobic surfaces have been reported by numerous authors5.

In most of these studies, the synthetic surface is fabricated by the photolithographic definition ofCorresponding authorsilicon wafers to create arrays of silicon posts with diameters ranging from 100nm6 to 100um7. Using photolithography to create superhydrophobic surfaces enables a wide variety of surface patterns to be formed with a high degree of precision.Despite the control afforded by photolithography, there are several limitations imposedby this process for the fabrication of useful, low-cost superhydrophobic components. Cost is one issue. Silicon wafer processing is expensive, especially considering the cost of tools and the time required to etch high aspect ratio, fine-pitch posts deep into the surface. Reliability of the surfaces is a second important factor. The silicon surface rapidly oxidizes to form SiO2 which is hydrophilic. Thus the nano/micro textured silicon surfaces must be treated with a coating to render the surface hydrophobic.

The silicon surface rapidly oxidizes to form SiO2 which is hydrophilic. Thus the nano/micro textured silicon surfaces must be treated with a coating to render the surface hydrophobic. Coatings include flurosilanes, fluoropolymers and plasma deposited fluorocarbon films. Such coatings on silicon have been found to exhibit adhesion loss with long term exposure to water8 causing the surface to lose its superhydrophobic properties. Finally, and perhaps most significantly, photolithography is inherently a planar process. The surface must be flat to insure the image remains in focus and maintain linewidth control. However, it is extremely challenging to fabricate channels and other complex, inherently three- dimensional surface shapes with a planar lithography approach.In this paper, we describe the use of solid freeform fabrication (SFF) technologies tofabricate superhydrophobic surfaces and 3D structures from hydrophobic polymers9.

The SFF approach address the three limitations imposed by lithographic processes as the SFF technologies are robust and offer a relatively low-cost route to fabricating superhydrophobic components. As the polymers used to fabricate the superhydrophobic structures are themselves hydrophobic, coatings are not required and adhesion issues are eliminated, rendering structures which are more reliable. Finally, three-dimensional parts, such as conduits, helices and membranes10 can be fabricated with commercial machines and readily available CAD modeling software.Ordered arrays of micro-textured features (>150 micron diameter) were built intoarbitrarily shaped three-dimensional structure including tubes and helices by using a commercially available polymer-jetting technique. The superhydrophobic properties, such as roll-off angle and water head pressure, were demonstrated to be governed by the same equations as for micro- and nano-scale superhydrophobic surfaces patterned on silicon.

In addition, we used a robotic dispensing system to build surfaces with silicone polymers which are not compatible with commercially available SFF systems.Experimental:Commercial 3D Printer: An InVision-HR 3-D printer from 3D Systems was used tofabricate parts. This machine uses multi-jet modeling technology combined with a UV-curable urethane-acrylic model material (VisiJet HR200) and a wax-based support material (VisiJet S100). After the build, the support material was removed in a forced-air convection oven at 70oC. Parts were designed using ProEngineer or 3dsMax software and exported as STL files to the InVision-HR.Robotic Dispenser: A robot (Janome-2203N) with +/- 10 micron repeatability was usedas an XYZ platform and trigger for a syringe controller (EFD Performus) and a UV lamp (Dymax BluWave 200) as shown schematically in Figure 1.

Two different materials were usedfor building the posts: A viscous room temperature vulcanizing (RTV) silicone (DAP aquarium sealant) and a UV curable silicone (Dymax Cure-Point 9440 AB-Mix). To control the rheological properties of the low viscosity 9440, 10 parts per hundred of fumed silica (Cab-o-Sil TS 530) was dispersed into the adhesive to provide thixotropy. The materials were carefully loaded into 10cc polyethylene syringe barrels, avoiding air bubble entrapment, and capped with either a 27 gauge (210 um i.d.) tapered tip with the RTV, or a 33 gauge (100um i.d.) chamfered steel tip with the UV curable adhesive. Both the syringe and UV light-guide were mounted parallel to each other on the robotic arm such that the adhesive drop was exposed to UV light only after the arm was translated by 18.5 mm after the dispense step was complete. Glass microscope slides were used as substrates and placed on top of an aluminum stage.

Glass microscope slides were used as substrates and placed on top of an aluminum stage. The dispense cycle program repeated the following sequence to create an array of posts:(cid:1) bring the arm to a predetermined dispense height above a substrate (cid:1) (cid:1) (cid:1) For single dispense posts: translate the syringe to the next dispense locationtrigger the syringe controller to dispense at a pre-set time and pressure lift the syringe normal to the surfaceor(cid:1) For multi-layer posts: translate the arm to position the UV light guide above the dispense location, trigger the UV lamp for a pre-set time and translate to the next dispense location.Figure 1: Schematic of the robotic dispensing system with UV lampExperimental Measurements: The height of a water column supported by a superhydrophobic sample was measured on samples covered with a 25mm x 25mm array of posts and fitting a square tube of the same material above the sample.

Water was added slowly to the chamber created by the tube and the sample base until the transformation from Cassie (non-wetted) to Wenzel (fully-wetted) state was observed visually. This transition was easily and reproducibly observed due to the change in refractive index between air and polymer. Contact angles were measured using a ramé-hart model 250 goniometer.Results and Discussion:Commercial 3D Printer: Parts made with the InVision-HR machine exhibited good imagefidelity and measured part dimensions agreed well with CAD specifications. Arrays of cylindrical (Figure 2a), square and square-based pyramidal posts (Figure 3) were successfully created.

Arrays of cylindrical (Figure 2a), square and square-based pyramidal posts (Figure 3) were successfully created. Droplets of water placed on these post arrays exhibited superhydrophobic behavior as shown in Figure 2b.Figure 2: (a) Cylindrical posts fabricated using a InVision 3-D printer measuring 175 micron diameter x 1300 microns tall; and (b) a 2mm diameter water droplet posed atop a similar surface demonstrating superhydrophobic behavior.Post arrays faithfully replicate the pitch and build height specified by the CAD model asillustrated for the case of square-based pyramidal shaped posts over the range from 300 to 500 micron pitch in Figure 3. Each square-based pyramidal post is designed to have a 200 x 200 micron base with a height of 1000 microns.

Each square-based pyramidal post is designed to have a 200 x 200 micron base with a height of 1000 microns. At 250 micron pitch, however, the posts overlap at the base reducing the relative post height by almost 40% from the nominal value of 1000 microns.Figure 3: Effect of post pitch on height and diameter of square-based pyramidal postsArrays of posts were integrated into the surface of more complex structures in CADmodels and successfully printed as monolithic objects. Tubes, conduits and helices were fabricated, either as building blocks (Fig 4a) which could be assembled into any arbitrary pathway (Fig 4b) or as contiguous objects, limited in size only by the constraints of the printing equipment (Figure 4c). Conical helix pathways, such as the one shown in Figure 4c, were fabricated both with and without arrays of posts integrated into the conduit surface.

Conical helix pathways, such as the one shown in Figure 4c, were fabricated both with and without arrays of posts integrated into the conduit surface. When no surface features were present, water droplets experienced sufficient drag that they remained stationary until a sufficient number of droplets coalesced such that the mass of water exceeded a threshold value allowing a large drop of water to slide partway down the path. When the conical helix conduit surface was formed with arrays of posts, however, water droplets rolled readily along the conduit due to the low drag at the water-surface interface.

When the conical helix conduit surface was formed with arrays of posts, however, water droplets rolled readily along the conduit due to the low drag at the water-surface interface. Releasing water droplets at a fixed frequency from a syringe formed a periodic stream of individual droplets with the same frequency.Figure 4 Superhdrophobic helical building block with connectors at each end (a) Four helical units assembled into a two-turn helix (b) A monolithic conical helix measuring 75mm x 75mm x 50mm tall (c).The behavior of water on superhydrophobic surfaces fabricated by 3D printing aregoverned by the same equations as those for superhydrophobic surfaces fabricated with nano- scale features11, which are based on a balance of surface tension and hydrostatic forces. This is illustrated in Figure 5 for the case of the fluid pressure that such structures can support.

This is illustrated in Figure 5 for the case of the fluid pressure that such structures can support. For water supported on round-headed posts, the fluid height can be calculated11:where γ is the surface tension, ρ is the density of the fluid, g is gravitational acceleration, D is the diameter of the post heads, L is the space between posts and θ is the advancing contact angle (113o for the HR200 model material). This relationship is plotted as the green curve using the diameter values shown in Figure 3. This plot has a similar shape to the experimental measured values (blue curve Fig 5, but diverges as the pitch decreases. This may result from the waterFigure 5: Effect of post pitch on the height of water that can be supporteddroplet penetrating partially into the post tips as has been observed for posts with <200nm diameter 11. Optimizing the fit of the theoretical curve to the experimental values requires increasing the post diameter by 20% (red curve in Fig 5).

Optimizing the fit of the theoretical curve to the experimental values requires increasing the post diameter by 20% (red curve in Fig 5). This would require the droplet descending approximately 130 microns into the posts, which is reasonable given their taper.The base for these experiments serves mainly to support the post arrays. However, sincethe droplet is suspended on (or near) the post tips, the base could equally well be made from a porous mesh, so long as the post pitch can be maintained. Porous superhydrophobic conduits have been constructed, as shown in Figure 6, such that aqueous fluids are supported whileFigure 6: (a) Porous superhydrophobic conduit 10mm in diameter(a); (b) Detailed view of (a) with posts measuring 150 um diameter x 1500 um tall; and (c) Schematic cross-section of pressurized superhydrophobic tube.permitting an unrestricted flow of gas through the tube.

In this way the droplet can interact with the gaseous atmosphere unrestricted by diffusion through a solid support base. By inserting this porous superhydrophobic tube inside a solid tube, the region between the solid and porous tubes may be pressurized with a gas. The gas pressure could be adjusted to maintain the liquid/gas interfacial pressure differential to within the value supported by the superhydrophobic micro- textured features.Robotic Dispensing System. The commercially available multi-jet printer easilyfabricated a variety of superhydrophobic parts. However this printer has certain limitations including material selection and resolution. To expand the range of materials that could be used to form superhydrophobic surfaces, we developed a robotic dispensing system with a UV lamp and light guide to enable in-situ curing of photosensitive materials.

By modifying the tipdiameter, dispense pressure and dispense time, both a high viscosity RTV silicone as well as a low viscosity UV curable adhesive were dispensed.RTV Silicone: Superhydrophobic surfaces were fabricated from arrays of RTV postswith pitches ranging from 450 to 700 microns. A single drop was used to fabricate each post. A ~5uL drop of water on a 500 pitch RTV post array is shown in Figure 7, which has an apparent contact angle >170 degrees. In contrast, the contact angle of water on a flat silicone sample is 110o The tilt angle required for a water droplet to roll off a RTV silicone superhydrophobic surface is 20o for volumes between 10-20 uL.

Droplets 8 uL and smaller required significantly higher angles (> 28o) as the diameter of the droplet becomes comparable to the pitch of the posts.Figure 7: Water droplet on an RTV silicon superhydrophobic surface with 500 um pitch postsThe shape of the post was dependent upon several factors including the syringe position relative to the substrate during dispensing (dispense height) as well as the direction of the syringe tip motion after dispensing was completed. Reducing the pitch further causes the bases of the posts to overlap whereas increasing the pitch above 700 pitch allows the water droplet to penetrate the post surface and wet the side walls and base.The effect of dispense height on post shape can be seen in Figure 8. When the syringe tipis close to the substrate, the RTV is forced to flow laterally, increasing the diameter of the post base (Fig 8a).

When the syringe tipis close to the substrate, the RTV is forced to flow laterally, increasing the diameter of the post base (Fig 8a). As the dispense height increases, lateral flow decreases, reducing the post base diameter and forming a post that has a more uniform diameter along its height (fig 8d). If the dispense height becomes too large the drop does not reproducibly touch the substrate resulting in missing posts. The optimal dispense height for uniform posts for a superhydrophobic surface is between 250um-500um.Figure 8: RTV silicon posts fabricated by dispensing through a 27 gauge tapered tip at height of: (a) 100 um; (b) 190 um; (c) 250 um; and (d) 300 um above a glass substrate.The shape of the posts can be controlled by adjusting the direction of the dispensing tipmotion after the dispense cycle is completed. For the posts in Figure 8, the tip was moved upwards, normal to the surface until the material disengaged from the tip.

For the posts in Figure 8, the tip was moved upwards, normal to the surface until the material disengaged from the tip. Angled posts wereintentionally made through a two step release process. Immediately after dispensing, the robot arm moved upwards 800 microns (normal to the surface), at which point the dispensed adhesive was still attached to the tip.

Immediately after dispensing, the robot arm moved upwards 800 microns (normal to the surface), at which point the dispensed adhesive was still attached to the tip. The arm then rises at a 63° angle during which time the adhesive detaches from the tip forming an asymmetric post in the direction of the tip motion as shown in Figure 9.Figure 9: Tilted RTV posts created by raising the dispenser at a 63o angle during release.These surfaces display a direction dependent roll off angle, where the roll off angle in the direction of the tilt is about 10° less than the roll off angle in the opposite direction.UV Curable Silicone: Superhydrophobic surfaces with pitches of 350,400,450 and 500 microns have been made by dispensing multiple layers of individual drops of the adhesive and curing after each dispense. Modifying the rheology of the Dymax adhesive was essential for controlling the diameter and height of the posts.

Modifying the rheology of the Dymax adhesive was essential for controlling the diameter and height of the posts. The viscosity of the as-received material was 1000 cps; dispensing this material would result in rapid flow across the substrate. By adding fumed silica, the low viscosity adhesive became thixotropic and so exhibited little flow before curing. With this rheology, further reduction in pitch should be possible, but would require the use of smaller diameter dispensing tips. The contact angle of water on a flat surface prepared from the Dymax UV curable silicone formulated with the thixotropic agent is 115°. This value is dramatically increased on a surface containing posts made of this material. Observed contact angles between this type of surface and a 5uL drop range between 157° and 169°.Similar to fabricating RTV posts, the height of the dispense tip above the previously deposited layer has a significant effect on the overall shape of the resulting post.

When the syringe tip is programmed to increment a relatively small amount between layers, the post spreads laterally causing the diameter of the post to increase as it rises above the surface (Fig 10a). Uniform post diameters were achieved by maintaining the correct dispense height for each layer as shown in Figure 10b.Figure 10: Effect of layer dispense height on the shape of UV cured silicone posts.Post diameter is strongly affected by the dispense duration as shown by the red curve in Figure 11. Shorter dispense times resulted in smaller post diameters, approaching the inside diameter of the tip. The diameter measured at the base of the post is larger than the overall post diameter due to wetting characteristics between uncured adhesive and glass substrate. For a dispense time of 200ms, the base spreading is 155% of the post diameter, while at 70ms it is minimized to be 100% (or equal to) the post diameter.

For a dispense time of 200ms, the base spreading is 155% of the post diameter, while at 70ms it is minimized to be 100% (or equal to) the post diameter. Decreasing dispense time significantly reduces the flow of excess material down the post as shown by the clearly delineated layers in Figure 12.Figure 11: Effect of dispense time in milliseconds on average post diameter in microns (red) and base spreading shown as the percent increase in diameter measured at the base relative to the post diameter (blue).Figure 12: UV cured silicone posts at 400 um pitch fabricated by dispensing 23 layers of adhesive. Posts measure 175 um in diameter and 800 um tall.Another parameter that affects post diameter, spreading and shape is the UV dose. Curing with UV hardens the silicone preventing flow and allowing for effective layer stacking. A dose of 100 J/cm2 was required to produce the 23 layer stack in Figure 12.

Curing with UV hardens the silicone preventing flow and allowing for effective layer stacking. A dose of 100 J/cm2 was required to produce the 23 layer stack in Figure 12. An insufficient dose resulted in short cone-shaped posts whereas additional cure times had no subsequent effect.Conclusions: Superhydrophobic surfaces were prepared on a wide variety of surface shapes using a commercially available 3D printer. Fabrication of large, complex polymer objects including conduits, helices and conical helix structures was achieved where the micro-textured surface is monolithic with the body. The superhydrophobic behavior, such as the water column height supported, is easily measured and can be described by the same equations as those used to describe superhydrophobic behavior on surfaces with nano-scale textural features.

By eliminating the need for hydrophobic coatings, these structures are expected to exhibit reliable superhydrophobic behavior.To expand upon the limited material set available with commercially available printers, a dispensing robot system was developed which accommodates a wide selection of materials with different rheological properties and cure requirements (room temperature & UV cured). Arrays of posts were formed either from the deposition of single droplets (high viscosity materials) or the build-up of multiple thin layers (thixotropic materials). In both cases, good control of post shape and feature dimensions were achieved.

In both cases, good control of post shape and feature dimensions were achieved. We anticipate that this approach will enable the fabrication of superhydrophobic components that will be useful for a variety of applications in the biological and separation science areas.Acknowledgements: This work was supported by IDA Ireland .and NYSTAR, the New York State Foundation for Science, Technology and Innovation.References1 Henoch, C.; Krupenkin, T.N.; Kolodner, P.; Taylor, J.A.; Hodes, M.S.; Lyons, A.M.; Peguero, C.; Breuer, K.S. Turbulent Drag Reduction Using Superhydrophobic Surfaces. 3rd AIAA Flow Control Conference, San Francisco, CA, June 5 – 8, 2006, 3192.2 Ahuja, A.; Taylor, J.A.; Lifton, V.; Sidorenko, A.A.; Salamon, T.R.; Lobaton, E.J.; Kolodner, P.; Krupenkin, T.N. Nanonails: A Simple Geometrical Approach to Electrically Tunable Superlyophobic Surfaces. Langmuir 2008, 24, 9-14.3 Neinhuis, C.; Barthlott, W.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFabricating Zirconia Parts with Organic Support Material by the Ceramic On-Demand Extrusion ProcessWenbin Li1, Amir Ghazanfari1, Devin McMillen1, Andrew Scherff1, Ming C. Leu1 and Gregory E. Hilmas21Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO, USA 2Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO, USAAbstractCeramic On-Demand Extrusion (CODE) is an extrusion-based additive manufacturing process recently developed for fabricating dense, functional ceramic components. This paper presents a further development of this process and focuses on fabricating 3 mol% yttria-stabilized zirconia (3YSZ) components that cannot be fabricated without using support structures.

This paper presents a further development of this process and focuses on fabricating 3 mol% yttria-stabilized zirconia (3YSZ) components that cannot be fabricated without using support structures. The 3YSZ paste is deposited through the main nozzle, and a polycaprolactone (PCL) pellet feedstock is melted and deposited through an auxiliary nozzle to build support structures. After a green part is printed and dried, the support structures are removed by heating the part to ~70 °C to melt the PCL. The part is then sintered at 1550 oC to achieve near theoretical density. The maximum angle of overhanging feature that can be fabricated without support was determined to be 60°.

The part is then sintered at 1550 oC to achieve near theoretical density. The maximum angle of overhanging feature that can be fabricated without support was determined to be 60°. Sample parts were fabricated and evaluated to demonstrate the effectiveness of the PCL support material and CODE’s capability to fabricate geometrically complex parts.IntroductionSeveral additive manufacturing (AM) processes have been developed for ceramics and glasses, including binder jetting [1,2], material extrusion [3-12], vat photopolymerization [13,14], powder bed fusion [15,16], and directed energy deposition [17-22], among others. Ceramic On- Demand Extrusion (CODE) [8-12] is a recently developed paste extrusion based AM process, which produces ceramic components with near theoretical density (>98%) after sintering. It deposits high solids loading (>50 vol%) aqueous ceramic pastes onto a substrate layer by layer at room temperature.

It deposits high solids loading (>50 vol%) aqueous ceramic pastes onto a substrate layer by layer at room temperature. Each deposited layer is partially solidified by uniform infrared radiation drying from above prior to initiation of a subsequent layer. At the same time, undesirable water evaporation from the sides of the part is prohibited by surrounding the part with a liquid [11]. This layered uniform radiation drying approach minimizes the moisture content gradient in the part during the fabrication process and thus enables CODE to produce crack-free ceramic parts (after sintering). The printed parts are then bulk-dried in a controlled environment with appropriate humidity, after which the green bodies are sintered to produce near-theoretical density parts [8- 10].Support structures are needed for the CODE process to produce geometrically complex parts.

In a previous study, an inorganic sacrificial support material, consisting of CaCO3 (calcium605carbonate), was used as the support material for fabricating Al 2O3 (aluminum oxide) parts. A two- step sintering process was developed for the support structure removal due to the favorable phase equilibrium between Al 2O3 and CaCO3 within the sintering temperature range. After the green part was fabricated, the first sintering step was performed at 1100 ℃. This step decomposed the support material into CaO (calcium oxide) and partially densified the main material. The part was then placed in water or an acid solution to dissolve the support structures. After the support material was removed, the second sintering step was performed at 1500 ℃ to obtain the dense ceramic part. Although successful, the multi-step sintering reduces the overall process efficiency.

Although successful, the multi-step sintering reduces the overall process efficiency. Also, in using an inorganic support material, co-firing is required to remove the support material, which introduces the risk of part cracking due to mismatch in the coefficient of thermal expansion (CTE) between the main material and the support material. Compared to an inorganic support material such as CaCO3, an organic support material can be removed from the part before sintering, and thus the undesirable co-firing process is avoided.The present paper describes the development of a technique for fabricating 3YSZ parts by the CODE process with the use of an organic sacrificial support material. A CODE system developed for this process was configured for concurrent deposition of two materials, with one auger extruder for depositing 3YSZ paste as the main material and a heating syringe extruder for melting and depositing an organic support material.

Polycaprolactone (PCL), which has a melting point of 60 ℃, was investigated for use as the support material. Sample parts including an H-shaped part and a spur gear were fabricated for demonstration purposes. The dimensions of these parts after sintering were measured and compared to their CAD models.CODE Process and Experimental SetupThe CODE process extrudes and deposits aqueous pastes onto a substrate to print layers sequentially. During printing, the level of a liquid medium is controlled to surround the part in order to prevent water evaporation from the sides of the part, thus avoiding moisture gradation from the part’s side surfaces to its core. Between the printing of each layer, infrared radiation is applied from the top to uniformly dry and partially solidify the layer. The layered uniform radiation drying prevents part cracking and warping.

The layered uniform radiation drying prevents part cracking and warping. More details of the CODE process are available from references [8-12].The CODE fabrication system [11] consists of a motion subsystem (gantry), an extrusion device mounted on the gantry and capable of extruding viscous ceramic pastes at controlled flowrates, an oil feeding device to regulate the oil level in the tank, and an infrared radiation heating device capable of moving the infrared source and providing on/off control. In the present work, the CODE system was configured for deposition of two materials. As shown in Figure 1, extruder A is an auger extruder for ceramic paste deposition [23], and extruder B is a heating syringe extruder for depositing a molten organic material by pneumatic force.606Figure 1.

An overall view (left) and a close-up view (right) of the CODE fabrication system configured for dual- extruder printing.Experimental ProcedureA commercially available zirconium oxide powder (TZ-3Y-E, Tosoh USA, Inc., Grove City, OH, USA) was selected as the raw material. Batches of ceramic paste were produced in 100 ml quantities and consisted of 52 vol% ceramic particles dispersed in distilled water, 30% ammonia ammonium hydroxide solution (NH4OH, Sigma Aldrich, St. Louis, MO, USA), R- C(O)OH (Dolapix CE 64, Zschimmer & Schwarz, Inc., Lahnstein, Rhineland-Palatinate, Germany) dispersant, polyethylene glycol (PEG 400, Sigma-Aldrich, St. Louis, MO, USA) and hydroxypropyl methyl cellulose (METHOCEL J 75 MS-N, DOW Chemical Company, Midland, MI, USA). The suspensions were mixed within 24 hours prior to part fabrication.

Louis, MO, USA) and hydroxypropyl methyl cellulose (METHOCEL J 75 MS-N, DOW Chemical Company, Midland, MI, USA). The suspensions were mixed within 24 hours prior to part fabrication. The solvent pH was adjusted drop-wise using ammonium hydroxide solution until an alkaline pH ≈ 9-10 was achieved, as measured by a pH meter (HI 2210, Hannah Instruments, Woonsocket, RI, USA). A zirconia slurry was prepared by combining the solids, ammoniated water, dispersant, and spherical 3 mm zirconia milling media (YSZ grinding media, Inframat Advanced Materials LLC, Manchester, CT, USA) in a HDPE jar, and was milled for 18-20 hours. PEG 400 and METHOCEL were then added and the contents were milled for an additional 3 hours to achieve homogeneity.

PEG 400 and METHOCEL were then added and the contents were milled for an additional 3 hours to achieve homogeneity. The final two added components increased the viscosity such that the slurry became a paste, which was then agitated on a vibratory table to remove entrapped air prior to transfer to a pneumatic syringe for deposition.Several organic materials were considered for potential use as the support material. Methyl cellulose solution was first examined. It was previously shown to work as a support material in the fabrication of Al 2O3 parts by the Freeze-Form Extrusion Fabrication process at -10 ℃ environment temperature [6]. However, this material turned out to be not printable in the CODE process since it could not solidify at room temperature until being dehydrated for hours. Candle wax and microcrystalline wax were then examined, but they are soluble in oil and became soft during the printing process.

Candle wax and microcrystalline wax were then examined, but they are soluble in oil and became soft during the printing process. Support materials for the Fused Deposition Modeling (FDM) process including PLA, ABS and PVA were not examined here since they would require about 200 ℃ printing temperature, which is not safe with the mineral oil used in the CODE process. Finally, PCL (polycaprolactone, Sigma Aldrich, St. Louis, MO, USA) was tested and turned out to be workable607at about 60 ℃. PCL pellets were loaded into the syringe of the heating extruder in the CODE system, and the heating temperature was regulated at 56 ℃, which is the lowest temperature to melt PCL pellets. An air pressure of 85 psi was applied to extrude the molten PCL.Results and DiscussionDetermination of maximum overhang angleThe maximum angle of overhanging features that can be printed without using support structures in the CODE process was determined for the 3YSZ paste.

Several wedge-shaped parts with different overhang angles were examined. The process parameters in this experiment are listed in Table 1. As shown in Figure 2, the overhanging feature started to deform when the test part with 65° overhang was attempted, and the part with 70° collapsed. Hence the maximum overhang angle for the 52 vol% solids loading 3YSZ paste in the CODE process was identified to be 60°. Note that the upper limit of the overhang angle may vary for different part material, part geometry and process parameters.Table 1. Process parameters for printing wedge-shaped test partsParameter Nozzle diameter Layer thickness Line spacing Printing speed Paste drying timeValue 610 µm 300 µm 600 µm 20 mm/s 25 s per layerFigure 2.

Wedge-shaped test parts attempted: (a) CAD models, (b) As-printed parts surrounded by oil in a tank, (c) Dried parts.Part fabrication608Several sample parts which required support structures during the CODE process were fabricated to validate the feasibility of using PCL as a support material and to evaluate the shrinkage of the fabricated parts due to sintering. An H-shaped part was chosen as the first sample geometry for evaluating the fabricated 3YSZ parts. Figure 3 shows the printing of an H-shaped part. Three H-shaped parts were fabricated using the process parameters in Table 1. The temperature of the heating extruder was 56 ℃ and the air pressure used for extruding PCL was 85 psi. The infill density for both the main material and the support material was 100%, i.e., solid infill. In addition to the H-shaped part, a spur gear with grooves was fabricated using the CODE process, as shown in Figure 4.

In addition to the H-shaped part, a spur gear with grooves was fabricated using the CODE process, as shown in Figure 4. Note that in Figure 4(b) red light was emitted from the infrared heater for drying each newly-deposited paste layer. The values of process parameters used for printing the spur gear were the same as those used for printing the H-shaped part.Figure 3. H-shaped part fabricated with using PCL as the support material: (a) CAD model, (b) & (c) Part being printed, (d) Part printed completely.609Figure 4. Printing of a spur gear sample part: (a) CAD model, (b) & (c) Part being printed, (d) Part printed completely.After each part was printed, the oil in the tank was drained and the part was removed, along with the substrate, from the tank for post processing.

Post processing included bulk drying for removing the remaining water content inside the part and also the oil on the part’s surface, followed by support material removal and part sintering.Bulk drying was accomplished by placing the CODE fabricated parts in an environmental chamber (LH-1.5, Associated Environmental Systems, Ayer, MA, USA), where the relative humidity and temperature were kept at 75% and 22 °C, for 12 h. The high humidity in the chamber provided a low drying rate, so that warpage and crack formation could be avoided. The substrate with the dried part was placed on a hot plate and heated to 70 ℃ to melt the PCL support structures. The ceramic green part was then removed from the substrate and sintered in a furnace (Deltech Inc., Denver, CO, USA), as shown in Figure 5. The furnace first heated the part to 500 ℃ and was held for 1 h to remove the binder content and residual PCL support material, and next ramped to 1550 ℃ and held for 1 h to sinter the 3YSZ part.

The furnace first heated the part to 500 ℃ and was held for 1 h to remove the binder content and residual PCL support material, and next ramped to 1550 ℃ and held for 1 h to sinter the 3YSZ part. Figure 6 shows the sintered spur gear.610Figure 5. Support material removal and sintering: (a) Substrate heated to melt PLC, (b) & (c) Part separated from substrate, (d) Parts placed inside a furnace for sintering.Figure 6. Final spur gear part after sintering.Part CharacterizationThe key dimensions of the sample parts (H-shaped part and spur gear) are given on the CAD models in Figures 3(a) and 4(a). These dimensions were measured on the sintered parts and compared to the values in the CAD models to evaluate part shrinkage.Measurements were taken from three H-shaped parts, with each dimension measured six times at random locations. The mean values and standard deviations of the six measurements are provided in Table 2 for each key dimension for each of the three samples.

The mean values and standard deviations of the six measurements are provided in Table 2 for each key dimension for each of the three samples. The overall mean values and average shrinkages (over the three samples) for the key dimensions compared to the CAD model dimensions are also given in Table 2. The key dimensions measured from the sintered spur gear and their means and standard deviations compared to the CAD model dimensions are given in Table 3. For the spur gear ten measurements were taken at random locations for each dimension.611Table 2. Dimensional measurements for the sintered H-shaped parts.MeasurementsSample 1Sample 2Sample 3AllMean Std. Dev. Shrinkage Mean Std. Dev. Shrinkage Mean Std. Dev.

Dimensional measurements for the sintered spur gear part.Measurements1 2 3 4 5 6 7 8 9 10 Mean Std. Dev. Shrinkage Max. Min.H1 (mm) 7.0 7.0 7.0 7.0 7.0 6.9 6.9 7.0 7.0 7.0 7.0 0.04 22.5% 7.0 6.9D1 (mm) 34.8 34.8 34.9 34.9 34.9 35.0 34.8 35.0 34.9 34.9 34.9 0.07 22.5% 35.0 34.8D2 (mm) 26.2 26.3 26.3 26.2 26.2 26.2 26.3 26.2 26.2 26.2 26.2 0.05 22.2% 26.3 26.2D3 (mm) 4.7 4.6 4.7 4.7 4.7 4.7 4.7 4.7 4.7 4.7 4.7 0.03 21.9% 4.7 4.6DiscussionFor the H-shaped part, the standard deviations of the measurements in Table 2 varied from 0.04 mm to 0.1 mm, and the shrinkages were in the range of 22.1% to 22.8%. The amounts of shrinkage of the spur gear after sintering given in Table 3 were in the range of 21.9% to 22.5%. For both parts, no significant anisotropic shrinkage was observed in both horizontal and vertical directions.

For both parts, no significant anisotropic shrinkage was observed in both horizontal and vertical directions. The diameters of the circular features of the spur gear, i.e., D1, D2 and D3 in Table 3, had standard deviations between 0.03 mm and 0.07 mm, corresponding to a range of 0.2% to 0.6%. This indicates good circularity in the fabricated spur gear.The above experiments and results validated PCL as a feasible support material for the CODE process. With the PCL support material, the green part can be separated from the substrate after being heated above 60 ℃, and the remaining PCL support material in the green part can be burned out in the furnace during the sintering process. However, there were issues observed when the molten PCL was deposited on top of a ceramic paste layer instead of the substrate. When PCL material was deposited onto a ceramic paste layer, it would warp and eventually detach from that paste layer after printing PCL for several layers.

When PCL material was deposited onto a ceramic paste layer, it would warp and eventually detach from that paste layer after printing PCL for several layers. The warped and detached support structure affected the dimensional accuracy of the printed part and imposed a high risk of damaging the part.612This is because when the deposited molten PCL on a ceramic paste layer started to deform due to thermal stresses while cooling down [24,25], the surface bonding provided by the ceramic paste was not strong enough to avert warping and detachment. A potential solution is using an organic material as the support material in the form of a paste, which is extruded and deposited without melting and hence avoid the warping and detachment. This idea will be investigated in a future study.ConclusionsA method of fabricating ceramic parts by the Ceramic On-Demand Extrusion (CODE) process with an organic sacrificial support material has been developed.

PCL (polycaprolactone) was identified as a feasible sacrificial material, which is melted and extruded to build support structures during the CODE process. The support structure can be removed from the part by heating the part above 60 ℃ to melt PCL and separate the green body from the support structure. The remaining PCL in the green body is then burned out during the part sintering process. Sample zirconia (3YSZ) parts including an H-shaped part and a spur gear were successfully fabricated by the CODE process using PCL as the support material. The dimensional shrinkages for the sintered parts compared to their CAD models ranged between 21.9% and 22.8%, and no significant anisotropic shrinkage in the horizontal and vertical directions were observed.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFabrication and characterization of Graphite/Nylon 12 composite via binder Jetting additive manufacturing processHadi Miyanaji1, Junaid Muhammad Akbar2, Li Yang11Department of Industrial Engineering, University of Louisville2Department of Mechanical Engineering, University of LouisvilleAbstractNowadays, graphite is used in many applications due to its unique combination of physicalproperties. Fabrication of graphite parts has been mostly restricted to traditional manufacturingprocesses (e.g. moldering), and limited works have been devoted to the feasibility of additivemanufacturing (AM) technology to produce graphite components. In the present study, thefeasibility of binder jetting additive manufacturing (BJ-AM) process in fabrication ofgraphite/nylon composites is investigated.

In the present study, thefeasibility of binder jetting additive manufacturing (BJ-AM) process in fabrication ofgraphite/nylon composites is investigated. The printability of the composite parts with varyinggraphite amount was experimentally examined through the adjustment of in-process parameters(e.g. saturation level and drying energy) and post-processing curing (e.g. curing time andtemperature). The efficiency of nylon as the liquid phase sintering agent was studied via themechanical property evaluation of the composites. In addition, the electrical properties of thegraphite/nylon composites were investigated in order to evaluate the effectiveness of themanufacturing method for graphite-based structures for potential functional applications.Keywords: Binder Jetting Process, Saturation, Composite, Liquid phase sintering1.

IntroductionAdditive Manufacturing (AM) technology is referred to as a collection of processes which createsa part geometry through adding materials layer by layer from a digital CAD file. AM allows forhighly complex parts to be made while wasting little to no material. One of the popular AMprocesses is the binder jetting additive manufacturing (BJ-AM) process. BJ-AM utilizes liquidbinder to selectively join powder from powder feedstock in order to create 3D geometries. For thefabrication of each layer, after a layer of powder material is spread on top of the powder bed, aprinthead selectively deposits certain amount of binder droplets at designated locations of thepowder bed. After binder deposition, the entire powder bed surface is subjected to external heatingapplied via a radiation heat source in order to partially cure the binder and ensure adequate593mechanical strength for subsequent printing processes.

After the drying/curing, a new layer ofpowder will be added to the powder bed, and the process repeats until the entire geometry is printed[1]. Upon completion of the printing process in order to increase the mechanical properties andstructural integrity of the green parts, the fabricated parts are further heated in an oven commonlyfor few hours at temperatures about 200 °C. After curing in the oven, the green parts possess onlylimited strength and often require additional post-processing such as sintering and infiltration inorder to achieve desired mechanical strength [2-4]. Due to the use of binder for geometry creation,BJ-AM technique possesses various advantages over other AM processes, such as the eliminationof thermally induced defects (e.g.

Due to the use of binder for geometry creation,BJ-AM technique possesses various advantages over other AM processes, such as the eliminationof thermally induced defects (e.g. distortion, unwanted grain growth, etc.) and broad compatibilitywith exotic materials such as ceramics and refractory metals [2, 5-8].Graphite is a commonly used material in many industries such as chemical industry, nuclearindustry, manufacturing industry, etc. It is also a good candidate for refractory materials (e.g.crucibles, ladles, and molds for containing molten metals) due to its high temperature stability andchemical inertness. Graphite is part of the carbon family and has several desired traits andinteresting properties. Most applications utilize unique characteristics of graphite, such as thermalresistance and electrical conductivity or thermal conductivity, self-lubrication, etc. combined withother functional materials as composite structures [9].

combined withother functional materials as composite structures [9]. The commonly used method for fabricationof graphite/polymer components is the molding of the papered graphite/polymer mixture toproduce the desired geometries. There exist different mixing techniques for manufacturingpolymer/graphite composites including in situ polymerization, solution compounding and meltblending [10, 11]. In in-situ polymerization method, the monomer is polymerized in the presenceof the graphite and as a result, the in-situ compounding technique offers stronger interactionsbetween the graphite and the polymeric phase. In solution compounding, the polymer is dissolvedin a solvent and graphite is dispersed in the resulting solution. In the melt bending technique,traditional mixing equipment such as extruder, internal mixer, and two-roll mill can be adopted forthe blending operations and are usually available in most compounding units.

Some of therestrictions of these processes includes high cost of mixture preparation, high cost of moldfabrication, part design limitations, etc.The desired properties of graphite coupled with complex geometries that cab be realized by meansof the BJ-AM could potentially benefit many industrial applications such as in electronics,computers, casting, electrodes and electrical contacts, etc. One of the challenges in fabrication of594graphite components using the AM process is the transformation of graphite from solid phase togaseous phase which might take place during sintering of the green parts at high temperatures andunder atmospheric pressure [12]. As a result, traditional sintering methods are not suitable to beemployed for the post-processing the graphite green parts.

As a result, traditional sintering methods are not suitable to beemployed for the post-processing the graphite green parts. In this study, in order to overcome thisissue, nylon-12 polymer powder was added to the graphite powder in varying volumetricpercentages in the attempt to enable liquid phase sintering in the printed components during post-processing. The added polymer powder is expected to melt and distribute through the graphitematrix under the capillary effects, and consequently increase the mechanical strength of the partsafter sintering. The effectiveness of nylon powder as the liquid phase sintering agent isexperimentally investigated in terms of mechanical strength (compressive strength) of the printedsamples. In addition, the electrical characteristics of the graphite/nylon composites are studied inorder to evaluate the effectiveness of the binder jetting process for graphite-based structures forpotential functional applications.2.

MethodologyGraphite powder provided by Superior Graphite company and nylon-12 powder material from 3DSystems company were used in varying proportions for preparing the graphite/nylon mixtures.Figure 1 shows the morphology of the powder materials used in this study. Although both powdersexhibit irregular particle forms, the nylon-12 particles appear to be shaped more spherically withno sharp edges and relatively consistent in size which attribute to good flowability.a)b)Figure 1. SEM images of a) Graphite powder b) Nylon-12 powder5954 different mixing ratios of graphite to polymer powder were set by volume. The powders werecarefully measured and mixed in a clean container before printing. The 4 different ratios ofpowders used in the study are shown below in Table 1.Table 1.

The powders werecarefully measured and mixed in a clean container before printing. The 4 different ratios ofpowders used in the study are shown below in Table 1.Table 1. Graphite-polymer powder volumetric ratios used for preparing the mixturesPowder mixture batch Graphite powder % Nylon-12 powder %10100225753505047525Ten cubes, each with dimension of 10 mm x 10 mm x 10 mm were printed per batch using theExOne M-Lab binder jetting printer. In order to achieve the desired mechanical strength of greenparts, different saturation levels (the liquid binder amount deposited for each layer of the parts)were examined while the remainder of the process parameters were kept constant during thefabrication process.

For all the builds, the in-process binder curing settings (curing power and time)referred to as energy output from the infrared heating lamp used to cure the binding agent duringthe printing process were set at 75% and 25 seconds for curing power and curing time, respectively,which were determined via preliminary process feasibility studies. Table 2 shows the processparameters used for printing of the parts.Table 2. Process parameters used for the fabricationParametersSaturation level (%)Power level (%) Drying time (s) Spread speed (mm/s)Examined level(s)70, 80, 9075251.5The thickness of each powder layer was fixed as 50 micrometers, therefore each of the 10 mmcubes consisted of 200 layers once fully printed. Moreover, through trial and error the powderspreading speed of 1.5 mm/sec was selected which resulted in uniformly spread layers. Figure 2shows the fabricated green parts before liquid phase sintering process.

Figure 2shows the fabricated green parts before liquid phase sintering process. After fabrication, all theparts were dimensioned using digital calipers (0.01 mm resolution) and weighed using a highresolution digital scale (0.0001g). The compression tests were performed on a Shimadzu micro-tensile testing machine using a 500 N load cell for the green parts and on a Instron 5569A with a5965 KN load cell for the sintered samples. In order to evaluate the electrical properties of thefabricated parts, Fluke 87-5 was used to assess the electrical resistivity of the samples fabricatedwith varying graphite percentage before and after post-processing.

In addition, the microstructureof the sintered samples was analyzed with the aid of ZEISS SUPRATM 35VP SEM equipment tostudy the effectiveness of liquid phase sintering in the micro scale.For the liquid phase sintering experiments, parts printed using each of the powder ratios weresintered at 180 °C for 2 hours and at 190 °C for 1 hour in an ExOne electric oven to evaluate theeffect of sintering time and temperature on integrity and mechanical strength of the samples. Toeliminate variation in sintering, all the samples that needed to be sintered at 180 °C were sinteredtogether in one batch and all the parts that needed to be sintered at 190 °C were sintered togetherin another batch. The sintering was performed in a SentroTech 6” High Temperature VacuumFurnace in atmospheric environment.Figure 2. The printed green parts with different saturation level and various mixtures of graphite/nylon-123.

The printed green parts with different saturation level and various mixtures of graphite/nylon-123. Result and discussionCompression strength of each specimen was calculated using the equation below.𝜎𝑐 =𝐹𝑐,𝑚𝑎𝑥 𝐴𝑐(1)597where 𝐹𝑐,𝑚𝑎𝑥 is the max compression force in Newton and 𝐴𝑐 is the cross sectional area of thespecimens in meter squared.The results of compression experimentations for green parts printed with different saturation levelsand varying graphite contents are shown in Figure 3. The figure shows that in general as saturationlevel increases, the compressive strength of the un-sintered parts increases. Such correlationbetween the saturation amount and the strength of the printed parts has been reported by variousresearchers for different combinations of binder and powder materials [5, 7, 13, 14].

Higher liquidbinder amount in a predefined envelope expands and strengthens the network of the particlesconnected by the binding agent and as a result, increases the mechanical strength of the printedcomponents. The results also showed that as graphite percentage increases, compressive strengthof the green parts exhibits a decreasing trend.700.0600.0) a p k ( h t g n e r t S n o i s s e r p m o C500.0400.0300.0200.070% Saturation80% Saturation90% Saturation100.00.0100% Nylon-1225% Graphite50% Graphite75% GraphiteFigure 3. Average compressive strength of green partsFigure 4 shows the average compressive strength comparison for parts sintered at 180 °C for 2hours. The figure exhibits similar trends as the un-sintered results from Figure 3.

The figure exhibits similar trends as the un-sintered results from Figure 3. However, themagnitudes of the compressive strength for the sintered parts are greater as expected.598100000) a p K ( h t g n e r t S n o i s s e r p m o C1000010001001070% Saturation80% Saturation90% Saturation1100% Nylon-1225% Graphite50% Graphite75% GraphiteFigure 4. Average compressive strength of the parts sintered at 180 °C for 2 hoursIn general, higher saturation leads to higher compressive strength even for the sintered parts in thatthe higher amount of binding agent in the microstructure of printed samples helps expanding theconnection areas between particles and consequently increases the mechanical properties of theparts. For the case of graphite/nylon mixture, since the post-processing temperatures examined(i.e.

For the case of graphite/nylon mixture, since the post-processing temperatures examined(i.e. 180 and 190 °C) are less than the temperature required for the binder burn-out from themicrostructure, sintering the parts would not eliminate the binder saturation effect on mechanicalcharacteristics of the fabricated components. At temperatures, higher than the melting point ofnylon-12 (i.e. 175 °C), the nylon material melts and disperses through the graphite matrix of theprinted parts as the liquid phase sintering agent. The degree of the dispersion and the connectednessof nylon-12 network which largely determines the strength of the sintered components dependsnot only on the sintering parameters (i.e. temperature and time of sintering) but also on the ratioof materials composing the mixture. With higher magnitude of nylon-12 in the mixture, dispersionof the liquid phase sintering agent in the microstructure becomes more uniform and sufficient asshown in Figure 5.

With higher magnitude of nylon-12 in the mixture, dispersionof the liquid phase sintering agent in the microstructure becomes more uniform and sufficient asshown in Figure 5. Therefore, such trend of increase in the mechanical properties of the sinteringparts is expected as amount of the nylon material in the mixture increases. Figure 6 shows themicrostructure of a part sintered at 180 °C for 2 hours with high magnification. The circled regionon the figure exhibits a close-up image of graphite/nylon-12 interaction zone with graphiteparticles surrounded by the molten nylon-12 materials.599a)b)Figure 5. SEM micrographs of specimens containing a) 25% graphite b) 50% graphiteFigure 6. SEM image of a part sintered at 180 °C for 2 hours- Graphite particles surrounded by melted nylon-12The average compressive strength of the parts processed at 190 °C for 1 hour is shown in Figure7.

Parts sintered at 190 °C for 1 hour exhibit similar compressive strength behavior as previouslydiscussed for other scenarios. The compressive strength increases with higher saturation levels anddecreases with higher graphite content.600100000) a p K ( h t g n e r t S n o i s s e r p m o C1000010001001070% Saturation80% Saturation90% Saturation1100% Nylon-1225% Graphite50% Graphite75% GraphiteFigure 7. Average compressive strength of the parts sintered at 190 °C for 1 hoursFigure 8 shows direct comparison between the two sintering sets in order to investigate theefficiency of the examined sintering levels.

From the figure, it can be concluded the two sinteringlevels, 180 °C for 2 hours and 190 °C for 1 hour, aren’t significantly different and, there isn’t anobvious pattern or correlation between sintering levels examined and average compressive strengthof parts.10000010000) a p K ( h t g n e r t S n o i s s e r p m o C1000100101180 C - 2 Hours190 C - 1 Hour25% Gr 70% Sat25% Gr 80% Sat25% Gr 90% Sat50% Gr 70% Sat50% Gr 80% Sat50% Gr 90% Sat75% Gr 70% Sat75% Gr 80% Sat75% Gr 90% SatFigure 8. Direct comparison between compressive strength of parts post-processed with both sintering parametersThe electrical resistivity of the parts printed at 90% saturation level before and after sinteringperformed at 180 °C for 2 hours is shown in Figure 9. Two vertical axes with different scales (KΩ601and Ω) were used for graphing the results due to the significant difference of the sample’s electricalresistivity before and after sintering.

Two vertical axes with different scales (KΩ601and Ω) were used for graphing the results due to the significant difference of the sample’s electricalresistivity before and after sintering. From the figure, as the graphite magnitude increases in themixture, the electrical resistivity significantly decreases for both states, before and after sintering.Also, it is observed that the sintering remarkably affects the electrical resistivity of the specimens.Such variation in electrical properties of the samples with post-processing can be partiallyattributed to the porosity reduction in the microstructure of the samples during sintering in whichthe molten nylon diffuses through the graphite particles and fill the empty spaces (porosity).

Dueto the lower electrical resistivity of the nylon-12, 1012 Ω, compared to that of air which isapproximately 1016 Ω, porosity reduction during the post-processing might facilitate the currentflow through the microstructure of the printed parts. Another possibility that might explain sucheffect of post processing is the rearrangement of the graphite particles during sintering due to theforces involved in the interaction between the molten nylon-12 and the particles including thecapillary forces.Figure 10 exhibits the electrical resistivity of the same batch of the samples sintered at 190 °C for1 hour.

The trends observed in the figure are similar to those of parts sintered at 180 °C for 2 hours.It appears that sintering the samples at lower temperatures for longer period of time (at 180 °C for2 hours) results in slightly better electrical characteristics of the samples likely owing to the moreefficient rearrangement of particles during longer sintering time period.301600)Ω K(2514001200)Ω(y t i v i t s i s e r20151000800y t i v i t s i s e rll a c i r t c e l E105600400200a c i r t c e El0025% Graphite50% Graphite Graphite magnitude (%)75% GraphiteResistivity before sintering (KΩ)Resistivity after sintering (Ω)Figure 9.

Electrical resistivity of the samples before and after sintering (at 180 °C for 2 hours)602301600)Ω K (2514001200)Ω(y t i v i t s i s e r20151000800y t i v i t s i s e rlla c i r t c e El10600400a c i r t c e El52000025% Graphite50% Graphite75% GraphiteGraphite magnitude (%)Resistivity before sintering (KΩ)Resistivity after sintering (Ω)Figure 10. Electrical resistivity of the samples before and after sintering (at 190 °C for 1 hour)ConclusionIn the present study, the BJ-AM process was applied to fabricate graphite/nylon-12 components.It was shown that the liquid phase sintering process during which nylon-12 melts and dispersesthrough graphite particles was effective in significantly strengthening the printed parts. It was alsofound that the higher saturation levels generally improved the compressive strength of printedparts.

It was alsofound that the higher saturation levels generally improved the compressive strength of printedparts. As graphite content increased, the compressive strength of the parts decreased. Within therange of sintering parameters examined, 180 °C for 2 hours and 190 °C for 1 hour, the processedparts performed similarly in terms of mechanical properties. Wider domain of sintering levelscould provide more insight into the effect of post-processing parameters on the quality andintegrity of the parts fabricated using BJ-AM process. Furthermore, it was shown that the electricalresistivity of the printed samples is significantly influenced by the graphite quantity in the mixtureas well as post-processing parameters.603AcknowledgementThe authors are grateful of the material support from Superior Graphite Co. Ltd (Chicago, IL). The authorsalso want to acknowledge the support from Rapid Prototyping Center (RPC) at University of Louisville.References[1] E.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFabrication and characterization of Ti6Al4V by selective electronbeam and laser hybrid melting Bin Zhou1,2,3, Jun Zhou1,2,3, Hongxin Li1,2,3, Feng Lin1,2,3, * 1Department of Mechanical Engineering, Tsinghua University, Beijing 100084, China 2Biomanufacturing and Rapid Forming Technology Key Laboratory of Beijing, China 3Key Laboratory for Advanced Materials Processing Technology, Ministry of Education of ChinaAbstractA hybrid process, which combines electron beam selective melting(EBSM) and selective laser melting(SLM), is proposed in this study. Laser is led into the vacuum chamber through the lens so that laser can be used to fabricate the metal powder at the same time with electron beam.

Laser is led into the vacuum chamber through the lens so that laser can be used to fabricate the metal powder at the same time with electron beam. In this study, Laser is used to pre-preheat the metal powder in order to preventing powder spreading and laser is also used to fabricate the contour of the parts both inside and outside. Electron beam is used to preheat the metal powder to the specified temperature and to fabricate the interior of the parts. It can be sure that through the hybrid process we can fabricate the parts with both better surface quality, higher precision and higher efficiency. Ti6Al4V samples were fabricated by selective electron beam and laser hybrid melting. The surface roughness of the parts was measured, the microstructures of the contour and interior were characterized using scanning electron microscopy (SEM).

The surface roughness of the parts was measured, the microstructures of the contour and interior were characterized using scanning electron microscopy (SEM). The results are that as-fabricated parts have better surface quality than the parts fabricated only using EBM process.Keywords: Electron beam selective melting; Selective laser melting; Hybrid; Ti6Al4VIntroductionSelective laser melting (SLM) and electron beam selective melting (EBSM) are powder-based additive manufacturing technology, which utilize a focused electron beam/laser beam to melt and fuse metal powder by rapid self-cooling, and produce parts layer by layer with high accuracy. So that both SLM and EBSM can be applied in impeller, blade, porous structure and medical implants manufacturing [1-2]. SLM process is with higher precision and better surface roughness, because laser spot diameter, layer thickness and powder diameter are smaller.

SLM process is with higher precision and better surface roughness, because laser spot diameter, layer thickness and powder diameter are smaller. But the material absorption rate of laser is low, so the powder layer of SLM is thin (20μm - 50μm), which led to the forming efficiency of SLM is low (about 20cm3/ h). EBSM has many unique characteristics such as high energy efficiency, high scan speed, and moderate operation cost. The material absorption rate for the electron beam is much higher(about 80%), the powder bed temperature of EBSM is much higher (about 600℃-1000℃)[3-5]. So EBSM process has a high forming efficiency (which can be up to 80cm3 / h). But the precision and surface roughness of EBSM is worse than that of SLM.

So EBSM process has a high forming efficiency (which can be up to 80cm3 / h). But the precision and surface roughness of EBSM is worse than that of SLM. The surface roughness of the EBM parts spread over a broad range but could often exceeds 30μm depending on various factors such as process parameters, material types, powder characteristics and surface orientations [6-9].1924It is difficult to obtain high precision and efficiency at the same time with single forming process. Jose Coronel [10] proposed a Multi-3D System which combines CNC router in additive manufacturing process, aimed to increase the accuracy of the printing parts. This process will reduce the forming efficiency and cause material waste. Li Yang [11] investigated the feasibility of utilizing electro polishing for the direct treatment of Ti6Al4V parts fabricated by EBM. This is the post-process and will cost material lost and distortion. H.

This is the post-process and will cost material lost and distortion. H. Schleifenbaum [12-16] developed a new prototype machine tool including a kW laser and a multi-beam system. The core of the parts can be fast manufactured with a large beam diameter while the skin is manufactured with a small beam diameter in order to assure the part`s accuracy and detail resolution. But high power laser beam causes serious powder spattering which has a negative effect in the process as a whole.In this study, a hybrid process was proposed, which combined electron beam selective melting(EBSM) and selective laser melting(SLM). The advantages of both EBSM and SLM can be obtained in the hybrid process with high precision and efficiency at the same time. Ti6Al4V samples were fabricated by the hybrid process.

The advantages of both EBSM and SLM can be obtained in the hybrid process with high precision and efficiency at the same time. Ti6Al4V samples were fabricated by the hybrid process. The surface roughness of the samples was measured, the microstructures of the contour and interior were characterized using scanning electron microscopy (SEM).Selective electron beam and laser hybrid melting processElectron beam selective melting(EBSM) process and selective laser melting(SLM) process are combined into a hybrid melting process. In the hybrid melting process, both the electron beam and laser can preheat and fabricate the powder at the same time or at interval. It shares one powder supply process and powder spread process and the whole system is in one vacuum chamber.Laser is used to fabricate the contour of the parts both inside and outside. Thus samples are fabricated with high precision and good surface roughness.

Thus samples are fabricated with high precision and good surface roughness. Electron beam is used to fabricate the interior of the parts with high scanning speed, which can increase the forming efficiency (as shown in Fig. 1).Fig. 1 EB and laser scanning section1925In the hybrid melting process, steps are shown in Fig. 2. First of all, electron beam is used to preheat the powder bed to the specified temperature; then, powder is supplied and spread; Next, laser is used to pre-preheat the metal powder in order to preventing powder spreading; and after that, electron beam is used to preheat the powder bed to the specified temperature; Subsequently, electron beam is used to fabricate the interior of the parts and laser is used to fabricate the contour of the parts.Fig.

2 Hybrid process flow diagramDevelopment of hybrid equipmentOn the basis of self-designed EBSM-250 electron beam selective melting system by Tsinghua lens of chamber wall,University, to vacuum chamber with dynamic focusing scanning system outside vacuum chamber (as shown in Fig. 3).laser wasledthroughtheFig. 3 (a) Hybrid forming system; (b) Schematic diagram of system1926The hybrid equipment characteristics are shown in Table 1.

3).laser wasledthroughtheFig. 3 (a) Hybrid forming system; (b) Schematic diagram of system1926The hybrid equipment characteristics are shown in Table 1. The maximum power of the electron beam is 3kW, and YLR-200 ytterbium fiber laser has maximum power of 200 W (IPG Laser GmbH, Germany).Table 1 The hybrid equipment characteristicsElectron beamLaserEffective build volume [mm×mm×mm] Power [W] Scanning speed [mm/s] Beam diameter [μm] Pressure [Pa] Powder bed temperature [℃]3000 0 ~ 20000 200 ~ 400200×200×200200 0 ~ 2000 100 ~ 2001×100 ~ 1×10-3 20 ~ 1000Since both the electron beam and laser scanning focusing system has their own coordinate systems, it is necessary to unified these two coordinate systems into the substrate coordinate system. Laser scanning head is placed on the top of the vacuum chamber next to the electron gun and has an angle (β=8°) rotation around the Y-axis direction.

Laser scanning head is placed on the top of the vacuum chamber next to the electron gun and has an angle (β=8°) rotation around the Y-axis direction. Also, the center of the laser scanning head has a distance with the center of electron gun (as shown in Fig. 4). Coordinate translation and rotation transformation is needed.Fig. 4 Three coordinate systemsRelative to the substrate coordinate system, the electron beam coordinate system has just translation transformation. But as to the laser coordinate system, it has both translation and rotation1927transformation.(x′y′z′) 1=(xyz) 1      10 0tx01 0ty00 1tz00 0 1       …………………… (1)From the translation transformation of formula (1), since t z=0, we can get:    ′ = + x x t ′ = + y ′ = z zyx ty………………………………………… (2)As to the laser coordinate system, two steps are needed.

0 11 0      ) 1) 1((′′′′′′zy=xyzx0 +′ xttxty0 +t′ ytz0 01 +t′ z0 001            100 00 α αcos −sin 00 α αsincos 0000 1           cosβ0 βsin 0010 0−sinβ0cos 0β000 1           γcos −sinγ0 0sincos0 0γγ0 00 01 0 0 1…………… (3)From the translation and rotation transformation of formula (3), since α=γ=0°, we can get: ′′ x    ′′z==((zz+ + t y+ + t ytt) ′′ y )′ y′ y⋅⋅( = + + (sinβ+ytycosβ−x t+ + xt′ yx t+ + xt′ xt′ x))⋅⋅cossinββ……………………… (4)For any point in the CAD file, it can be transformed into both the electron beam and lasercontroller by formula (3) and formula (4).In the hybrid melting process, since the forming process is carried out under vacuum, the evaporation of volatile material such as Al is much greater and the lenses will be steamed with metal vapor.

After a long time working, the laser power will be absorbed by the lens and eventually the lens will be broken. As a result, the hybrid process is seriously affected.Anti-evaporation system was designed to prevent the lenses (as shown in Fig. 5). In front of the lens, a second lens was placed, which was bigger than the behind one. The second lens could rotate round the axis. The laser working area which laser passes through to the substrate, was in one side and the brush was in the other side. As-designed system could allow lens working with cleaning at the same time and the brush would not disturb the process.1928     Fig. 5 Anti-evaporation systemHybrid melting experimentAtomized spherical Ti6Al4V powder (shown in Fig. 6) were used in hybrid melting process. The powder particle size was in the range of 13~55μm, with an average of 32μm, which is obtained by laser particle size analyzer Hydro 2000MU (A). The powder size distribution is shown in Fig.

The powder particle size was in the range of 13~55μm, with an average of 32μm, which is obtained by laser particle size analyzer Hydro 2000MU (A). The powder size distribution is shown in Fig. 7 and the composition is shown in Table 2.Element wt%Table 2 Ti6Al4V powder composition Ti BalanceAl 6.05V 3.94C 0.02N 0.02O 0.087Fig. 6 SEM image of Ti6Al4V powders1929Fig. 7 Ti6Al4V powder size distributionThree Ti6Al4V samples were fabricated with the electron beam and laser (as shown in Fig. 8). The size of the samples were 20mm×20mm. The hybrid process was shown in Fig. 8(b), which used laser scanning the contour and EB scanning interior of the sample. To be comparison, two samples were fabricated which respectively used laser (Fig. 8(a)) and EB (Fig. 8(c)) scanning both the contour and interior. The powder layer thickness is 0.03mm, and 100 layers were fabricated for each three samples. The laser power was 100W, with scanning speed 200 mm/s and hatching space 0.05mm.

The powder layer thickness is 0.03mm, and 100 layers were fabricated for each three samples. The laser power was 100W, with scanning speed 200 mm/s and hatching space 0.05mm. As for EB, the power is 180W, and scanning speed is 250 mm/s with hatching space 0.1mm. The pressure was 2×10-3Pa, and the substrate was made of Ti6Al4V with thickness of 10mm. The powder bed temperature is controlled to 700-800℃ (as shown in Table 3).Fig. 8 (a) Laser scanning contour and interior; (b) Laser scanning contour and EB scanning interior; (c) EB scanning contour and interior;1930Table 3 Ti6Al4V fabrication parametersPower [W] Scanning speed [mm/s] Hatching space [mm] Layer thickness [mm] Pressure [Pa] Powder bed temperature [℃]Electron beam 180 250 0.10.03 2×10-3 700 ~ 800Laser 100 200 0.05The morphologies and microstructure of the Ti6Al4V samples were analyzed using a scanning electron microscope. The defect and pore were measured through the analysis of SEM images.

The defect and pore were measured through the analysis of SEM images. The samples surfaces of samples were etched with Kroll's reagent: 1% HNO3, 2% HF and 97% of distilled water. The density of the Ti6Al4V parts were measured by the method of drainage, using digital machine (ADVENTURER, AR423DCN) with an accuracy of 0.001 g. The surface roughness was measured by Surface topography measuring instrument (TalySurf Intra, Britain).The surface roughness results are shown in Table 4. The surface roughness of the side is mainly influenced by electron beam or laser parameters, powder size, layer thickness and powder bed temperature. In this study, three samples were fabricated in the same condition which shared the same powder size, layer thickness and powder bed temperature. So the differences of the side surface roughness of the samples result in the differences of the power parameters.

So the differences of the side surface roughness of the samples result in the differences of the power parameters. From the measurement results, the side surface roughness of the samples with laser scanning contour and EB scanning interior (11.53μm) is better than the samples with EB scanning contour and interior (17.30μm) and a little worse than the samples with laser scanning contour and interior (9.75μm). Because the laser beam spot is smaller (100-200μm) than the electron beam (200-400μm), the side surface roughness of the samples fabricated by hybrid melting process increases. And because of the higher absorption rate of electron beam, temperature of the molten pool of the hybrid melting fabricated samples is higher than that of laser fabricated one. This may cause the side surface roughness a little worse. As to the top, the surface roughness is mainly influenced by electron beam or laser parameters. The input energy density is higher; the surface roughness is better.

As to the top, the surface roughness is mainly influenced by electron beam or laser parameters. The input energy density is higher; the surface roughness is better. Because of the high absorption rate of electron beam, the input energy density of electron beam is much higher than laser. As a result, the top surface roughness of EB is better than that of laser. Almost of the top areas of hybrid melting fabricated samples are scanned by electron beam, so the top surface roughness is almost the same with electron beam fabricated one.Surface roughness R a[μm]Table 4 Ti6Al4V surface roughness (b) Laser scanning contour and EB scanning interior(a) Laser scanning contour and interior(c) EB scanning contour and interiorTop14.702.283.14Side9.7511.5317.301931The microstructures of hybrid melting fabricated Ti6Al4V are shown in Fig. 9. The interior fabricated by electron beam is dense with almost no pores and defects (Fig. 9a and Fig. 9c).

9. The interior fabricated by electron beam is dense with almost no pores and defects (Fig. 9a and Fig. 9c). While, the contour fabricated by laser has many visible pores and defects (Fig. 9a and Fig. 9b). The laser power in not enough to melt the powder densely, which is needed to be increased. In the hybrid melting process, powder is fabricated layer by layer. The molten pool has high temperature and fast cooling speed, and the solidified part is repeatedly heated and cooled, which results in the formation and transformation of organization with the forming process. When metal molten pool solidifies, the temperature gradient along the building direction is much larger than the other direction, So the crystal grains grow in columnar crystal β phase along the building direction. Subsequently, with the rapid cooling of molten pool, β phase is transformed into metastable fine acicular martensite without diffusion.

Subsequently, with the rapid cooling of molten pool, β phase is transformed into metastable fine acicular martensite without diffusion. Then because of the effect of reheating, metastable fine acicular martensite breaks down, forming elongated and coarse lamellar α+β phase. The final microstructure is mainly composed of fine acicular martensite, elongated lamellar α phase and coarse lamellar α phase. Due to the material absorption rate of electron beam is much higher than that of laser, and the molten pool temperature is higher, so in the microstructure of electron beam fabricated Ti6Al4V, α phase is much coarser in the basket-weave microstructure (Fig. 9d-e).Fig. 9 SEM images of hybrid melting fabricated Ti6Al4V1932ConclusionsThe electron beam and laser hybrid melting process was proposed and the hybrid forming equipment was developed.

9d-e).Fig. 9 SEM images of hybrid melting fabricated Ti6Al4V1932ConclusionsThe electron beam and laser hybrid melting process was proposed and the hybrid forming equipment was developed. Ti6Al4V samples fabricated by electron beam and laser hybrid melting process, which used laser scanning the contour and electron beam scanning the interior, have better surface roughness than electron beam fabricated samples. And the hybrid process has a much higher forming efficiency than SLM process or EBSM process with post-treatments.AcknowledgementThis work was supported by ‘Suzhou-Tsinghua special innovation leading action’ program.References [1]. Ge Wenjun, Guo Chao, Lin Feng, Microstructures of Components Synthesized via Electron Beam Selective Melting Using Blended Pre-Alloyed Powders of Ti6Al4V and Ti45Al7Nb, Rare Metal Materials and Engineering, 44 (2015): 2623-2627[2]. Dylan Agius, Kyriakos I.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFabrication and Control of a Microheater Array for Microheater Array Powder SinteringNicholas Holt, Lucas Galvan Marques, Austin Van Horn, Mahsa Montazeri, Wenchao Zhou The AM3 Lab, Department of Mechanical Engineering, University of Arkansas at Fayetteville Fayetteville, AR, United States of America; email: zhouw@uark.eduABSTRACT Microheater Array Powder Sintering (MAPS) is a novel additive manufacturing process that uses a microheater array to replace the laser of selective laser sintering as the energy source. Most of the previous research on microheaters are for applications in gas sensing or inkjet printing. The operation temperature and response time of the microheater array are critical for the choice of sintering materials and printing speed for the MAPS process.

The operation temperature and response time of the microheater array are critical for the choice of sintering materials and printing speed for the MAPS process. In this paper, we present the fabrication, packaging, and control of a platinum microheater array that has a target operation temperature of 400°C and a response time of ~1 millisecond. First, we will present the fabrication process of a microheater array. The fabricated microheater array is then packaged for easy control and to serve as the printhead of the MAPS process. A PID controller is designed to control the temperature response of the microheater. Finally, the effectiveness of the controller is evaluated. Results show the fabricated microheater array satisfies the design requirements for the MAPS process.Key words: Additive manufacturing, MAPS, microheaters, MEMS, PID controls1.

Results show the fabricated microheater array satisfies the design requirements for the MAPS process.Key words: Additive manufacturing, MAPS, microheaters, MEMS, PID controls1. INTRODUCTIONMicroheater Array Powder Sintering (MAPS) is an emerging additive manufacturing (AM) technology which selectively sinters powder particles with a microheater array. In contrast to selective laser sintering (SLS) which uses an expensive laser to scan a powder bed pointwise to fuse particles, MAPS uses an array of microheaters as an energy source to deliver a heat pattern by bringing the microheater array in close proximity to a powder bed to sinter powder particles in a layer-wise fashion as illustrated in Figure 1 [1] [2].

MAPS promises significant increase in printing speed by increasing the number of microheaters and great energy savings because a microheater typically operates with a power of ~1 watt (compared to a laser usually around ~100 watts in SLS).To maximize the potential of the MAPS technology, the microheater array needs to satisfy several requirements: 1) Each microheater element in the microheater array needs to be able to deliver sufficient energy to sinter powder particles at a speed on par with that of a laser; 2) Each microheater element needs to be individually operated at a timescale (i.e., the timescale of the heating and cooling cycle) shorter than the sintering timescale; 3). The microheater array needs to be fabricated in a cost efficient manner and be scalable to a large number of microheater elements at minimal additional cost; 4).

The microheater array needs to be fabricated in a cost efficient manner and be scalable to a large number of microheater elements at minimal additional cost; 4). The packaging of the microheater array should be designed such that nothing protrudes from the microheater surface to avoid interference with the MAPS process;13145). The temperature of the microheater needs to be precisely controlled and the temperature distribution in the microheater needs to be as even as possible for consistent sintering quality.Figure 1. Overview of the MAPS printing process: the powder particles underneath the selectively activated microheaters are sintered as the printhead moves over; the through-vias and the gold lead are to connect the microheaters to external control circuit.As will be revealed in our literature review in the next section, microheaters have been typically used in gas sensing and thermal inkjet printing applications.

No research has been reported on designing, fabricating, packaging, and controlling a microheater array that can satisfy the requirements for the MAPS process. Therefore, in this paper, we present a study on a microheater printhead for the MAPS process, which includes the fabrication, packaging, and control of a microheater array. We first developed a MEMS fabrication procedure for a 2 by 2 microheater array, which demonstrates the scalability of this fabrication procedure to larger number of microheater elements. Then we developed a Quad Flat No-leads (QFN) packaging for the microheater array to connect it to the external control circuit. A control circuit and PID control loop are then developed to control the microheater temperature. The entire printhead is then tested and results show that it satisfies the requirements for the MAPS process.This paper is organized as following. Section 2 provides a literature review on the fabrication, packaging, and controlling of microheaters.

Section 2 provides a literature review on the fabrication, packaging, and controlling of microheaters. In section 3, a fabrication procedure of a microheater array is developed. Section 4 presents a QFN packaging design of the microheater array. The control of the microheater array is presented in section 5. Conclusions and future work are presented in section 6.2. LITERATURE REVIEWA microheater is a resistive heater with a resistive filament miniaturized to ~100nm to ~100µm with MEMS fabrication techniques. The filaments are generally made of heating materials such as titanium, platinum, molybdenum, tungsten, or polysilicon, to name a few [3] [4] [5] [6] [7]. Characteristics of these thin film microheaters are low power consumption, fast response, good heat confinement, good mechanical stability, and good fabrication yield [3]. Applications for these1315microheaters include micro-ignition for micro-propulsion systems and micro-explosive boiling [3].

Applications for these1315microheaters include micro-ignition for micro-propulsion systems and micro-explosive boiling [3]. Microheaters have been used in both 2D and 3D printing systems such as inkjet printing [8] [9], thermal printing commonly used in a point-of-sale (POS) printer [10], and selective heat sintering (SHS) [11]. In these printing technologies, the microheater usually operates at around 300 °C, which is not sufficient for MAPS.Microheaters are generally built on a substrate of glass or silicon. Glass is a good choice due to its low thermal conductivity. Silicon has also been used despite its high thermal conductivity due to its ease of processing with MEMS fabrication. Silicon underneath the microheater can be etched away to leave the microheater on a thin membrane of dielectric material to increase power efficiency [12].

Silicon underneath the microheater can be etched away to leave the microheater on a thin membrane of dielectric material to increase power efficiency [12]. The typical process flow to fabricating a heater using MEMS techniques is to grow a dielectric layer, use photolithography to pattern the substrate then deposit the microheater material and conductive leads using sputtering or e-beam evaporation [13] [14]. Some important considerations in designing microheaters are the heat transfer characteristics, geometry, and thermal response time. Microheaters transfer heat in three modes: conduction, convection and radiation [15]. Of these three modes, the conduction and convection are significant while radiation is insignificant at temperatures less than ~700°C for Ti or Pt microheaters [10] [15] [16]. Numerical simulations have been used to select materials by evaluating maximum temperature and power savings with different insulating layers [4].

Numerical simulations have been used to select materials by evaluating maximum temperature and power savings with different insulating layers [4]. Geometries can also be optimized for power savings, reduced stress profiles and even heat distribution [17] [18] [19]. Even heat distribution was particularly critical for MAPS to ensure even heating of the substrate. In our prior research, we have optimized the microheater geometry to improve evenness in heating and reduce thermal stresses using particle swarm optimization [2], but no microheater has been fabricated. Thermal response time is critical to the operation of a microheater, particularly in the MAPS application.

Thermal response time is critical to the operation of a microheater, particularly in the MAPS application. In previous studies, a thermal response time of 2ms for a temperature of 600°C was achieved with tungsten microheaters with a power consumption of 12mW [6] and a thermal response time of 1ms was achieved with a Pt/Ti heater reaching a temperature of 400°C using only 9mW of power [20], but they are not optimized for even temperature distribution.Packaging is a container for a MEMS device. Materials for packaging are chosen to withstand the operating conditions, which could be high temperature operation, high pressure, chemical resistance, mechanical and thermal shock and vibration [21] [22]. Metals, ceramics, silicon and plastics are all commonly used for microheater packaging. Metals are good for their robustness ease of assembly, mechanical integrity and chemical inertness in harsh environments [23].

Metals are good for their robustness ease of assembly, mechanical integrity and chemical inertness in harsh environments [23]. Ceramics are popular due to the ease in shaping and their material properties such as electrically insulating, hermetic sealing, thermal conductivity and chemical inertness [23]. However, very few research has been reported on microheater packaging.A microheater is essentially a very small resistance heater, as such it operates by passing an electric current across the filament to generate heat. Since the response time of a microheater is generally very fast, a sophisticated feedback system is required to control the temperature. MAPS requires a feedback system which allows a 1ms thermal response time. Proportional-integral- derivative (PID) controls have been effectively used to control the temperature of a microheater [24] [25].

Proportional-integral- derivative (PID) controls have been effectively used to control the temperature of a microheater [24] [25]. The feedback control for a microheater relies on a phenomenon in which the resistance of a conductor changes with its temperature; thus, the average temperature of some conductor may be determined through its change in resistance [26]. The temperature of the microheater is usually1316read through the resistance change of an additional metal filament which is near the microheater, rather than by reading the resistance of the microheater [25] [26], although the resistance change of the microheater has been successfully used for low temperature applications at 200°C [24]; MAPS requires the microheater to operate at a high temperature of at least 400°C and does not have room for additional temperature sensors.Although different aspects of microheater design and fabrication have been studied, the literature on microheater packaging and control is scarce.

A complete solution to a microheater printhead for the MAPS process does not exist in the literature. In the following sections, we will present the fabrication, packaging, and controlling of the microheater printhead for MAPS.3. FABRICATIONThis section is built on an optimized microheater design in our prior research [2], which used particle swarm optimization to minimize unevenness in temperature distribution and the thermal stress in the microheater. The heater geometry design is shown in Figure 2(a). In this section, we will present a fabrication procedure of a microheater array using this geometry design.3.1. Microheater FabricationFor the MAPS process, a microheater with the following criteria was desired: a maximum temperature of 400°C with a response time of 1ms must be achieved.

Microheater FabricationFor the MAPS process, a microheater with the following criteria was desired: a maximum temperature of 400°C with a response time of 1ms must be achieved. Platinum was chosen as the filament material due to its oxidation resistance and ability to reach the target temperature, and the geometry of the heater was chosen to decrease stress and evenly distribute heat across the microheater [1]. Schott’s AF 32 eco glass wafer was chosen as a substrate due to its low thermal conductivity and high glass transition temperature of 717°C [1] [27]. The designed heater was scaled to a 2x2 array, as can be seen in Figure 2.Figure 2. (a) Design of a microheater array with evaporated gold leads; (b) 3D Model of microheater die with leads evaporated onto the sides.

(a) Design of a microheater array with evaporated gold leads; (b) 3D Model of microheater die with leads evaporated onto the sides. This enables electrical connections to the side of the die, allowing the heaters to approach close proximity to a substrate.One critical design aspect of the microheater die used for MAPS is that the microheater must be able to approach the substrate to a gap of only a few microns. This makes it necessary to eliminate any wire bonding or other electrical connections that may protrude out of the surface of the microheater die. One approach is to connect the microheater to the external control circuit from the other side of the die, which can be done using a glass wafer with through-vias as illustrated in Figure 1, such as that offered by Schott [28], which can be an ideal solution for mass production.

Since we required only small batches of microheater dies, we have chosen to evaporate conductive copper leads along the sides of the die to enable electrical connections to the other side of the die as shown in Figure 2(b).1317techniques. Photolithography followed by e-beam evaporation were used to deposit the microheater filaments and conductive leads. A more detailed description of the fabrication process is described below and shown in Figure 3:The die described above was manufactured using MEMS fabricationFigure 3 – 2D and 3D representations of the fabrication steps of the microheater array: (a). the glass wafer is coated with a layer of photoresist; (b). the photoresist is exposed using the microheater mask and etched; (c). an adhesion layer of Ti followed by Pt filaments are deposited; (d). the photoresist is removed using acetone; (e). an adhesion layer of Ti followed by Au leads are deposited using the same procedure in steps (a) – (d).1.

the photoresist is removed using acetone; (e). an adhesion layer of Ti followed by Au leads are deposited using the same procedure in steps (a) – (d).1. A 500µm thick AF 32 eco glass wafer from Schott Inc. was used. 2. The wafer was cleaned using acetone and isopropyl alcohol, then rinsed with DI water and dried.3. AZ 5214-E photoresist was spin coated to a thickness of 2µm and prebaked. 4. UV photolithography was performed using the microheater mask, followed by post- baking and a second blanket exposure.5. The photoresist was etched using AZ MIF 300 developer. 6. E-Beam evaporation was used to deposit 100Å of titanium (adhesion layer) followed by 1500Å of platinum (heater filaments).7. Lift-off of the photoresist and metal layer was achieved using acetone in an ultrasonic bath.8. A second photolithography stage identical to that previously mentioned was used to pattern the wafer using the leads mask.9.

A second photolithography stage identical to that previously mentioned was used to pattern the wafer using the leads mask.9. E-Beam evaporation was used to deposit 100Å of titanium followed by 1000Å of gold. 10. Lift-off of the photoresist and metal layer was achieved using acetone in an ultrasonic bath.11. The dies were cut using a dicing saw such that the gold leads extended to the edge of the die.After the heaters are fabricated onto the glass substrate, the wafer is diced into the individual heater dies, consisting of four microheaters each with two leads as shown in Figure 3. After dicing, it is possible to fabricate the extended leads, which are deposited via e-beam evaporation. The1318leads consist of an adhesion layer of titanium, followed by a layer of copper and finally a layer of tin, which is included to protect the copper from oxidation. The extended leads are fabricated onto these dies using the following steps, as shown in Figure 4:1.

The extended leads are fabricated onto these dies using the following steps, as shown in Figure 4:1. Kapton tape was used to mask the sides of the dies in between each of the leads and to mask the top of the die where the microheaters are to avoid shorts between the leads. 2. The dies were put into a fixture to tilt them at a 45° angle such that both the side and top of the dies are exposed during evaporation.3. E-beam evaporation was used to deposit 100Å of titanium followed by 1000Å of copper and finally another 100Å of tin to extend the leads to the sides of the die.Figure 4 – Fabrication steps to extend leads. a. thin strips of Kapton tape are placed in between the leads and on top of the microheater array, b. the microheater dies are fixed at a 45° angle, c. extended copper leads are evaporated onto the top and sides of the microheater die.3.2. Microheater Packaging The packaging of the microheater is an integral part to the success of MAPS.

extended copper leads are evaporated onto the top and sides of the microheater die.3.2. Microheater Packaging The packaging of the microheater is an integral part to the success of MAPS. The packaging must accomplish the following tasks:1. Provide a fixture of the heater die to the rest of the printer assembly; 2. Provide a method of electrical connection to the heaters; 3. Allow the microheaters to approach a flat substrate to a distance of ~2µm.To satisfy these requirements, a basic QFN package was devised for the microheater die. The package is a two-layer printed circuit board (PCB) with an FR4 substrate, purple mask over bare copper, and ENIG finish.

The package is a two-layer printed circuit board (PCB) with an FR4 substrate, purple mask over bare copper, and ENIG finish. The package design can be seen in Figure 5.Figure 5 - Top (left) and bottom (right) views of the packaging design for the microheater dieThe extended side leads of the microheater die are bonded directly onto the conductive pads on the packaging using a silver epoxy – this allows for both electrical connection and a method of1319attachment between the die and the packaging. EPO-Tek® P1011 silver epoxy [29] was used to bond the heater die to the packaging using the following steps:1. The microheater die is attached to the package using Kapton tape such that the leads on the die are aligned with the contact pads of the package.2. The epoxy is manually applied using a needle such that it connects the individual leads to the contact pads (soldering can be used to replace epoxy in a production scenario).3.

The epoxy is manually applied using a needle such that it connects the individual leads to the contact pads (soldering can be used to replace epoxy in a production scenario).3. A weight is placed on top of the die to ensure it stays in place, and the package is loaded into a vacuum oven to bake at 150°C for one hour to cure the epoxy.4. The package is removed from the vacuum oven and the Kapton tape is removed. The finished packaging of the microheater die is shown in Figure 6.Figure 6 - Finished microheater packaging with electrical connections and attachment accomplished with conductive epoxy4. MICROHEATER TEMPERATURE CONTROLFor MAPS, it is important to have precise control of the temperature of the microheater. Ideally, the microheater should ramp to its target temperature of 400°C in 1ms and be maintained at that temperature until sintering is complete.

Ideally, the microheater should ramp to its target temperature of 400°C in 1ms and be maintained at that temperature until sintering is complete. It is also critical that the temperature of the microheater does not too far exceed the target temperature or else there is a risk of damaging the microheater. Because of these requirements, a sophisticated feedback control is needed.The microheater temperature is controlled by a PID controller, a closed-loop control based on the temperature-resistivity relation of the microheater. Figure 7 illustrates the control loop of the microheater temperature. All data is processed by an 8-bit microcontroller board Arduino Mega. Arduino Mega generates an 8-bit data output which is converted to an analog voltage between 0 and 10 volts by a Digital to Analog Converter (DAC). The signal goes through the power amplifier to ensure that enough power will be delivered to the Wheatstone bridge where the microheater is connected.

The signal goes through the power amplifier to ensure that enough power will be delivered to the Wheatstone bridge where the microheater is connected. The voltage difference in the Wheatstone bridge between the Point A and Point B (Figure 8) is converted by the Analog to Digital Converter (ADC) and sent back to the Arduino Mega where the resistance of the Microheater is calculated.

Based on the temperature-resistance relation, the temperature of the microheater is calculated and the control voltage output is updated to keep the microheater temperature around the target temperature.1320Figure 7 - Block diagram of the microheater control unitFigure 8 - Microheater connected in the Wheatstone bridgeThe microheater resistance Rm is calculated using three known resistors in the Wheatstonebridge as shown in the formula below:𝐴𝐴𝐴𝐴𝐴𝐴 𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉 𝐴𝐴𝐴𝐴𝐴𝐴 𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉⎛𝑅𝑅2 ∗ 𝑅𝑅3 + �𝑅𝑅3 ∗ (𝑅𝑅1 + 𝑅𝑅2) ∗ ⎜𝑅𝑅𝑚𝑚 =𝐴𝐴𝐴𝐴𝐴𝐴 𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉 𝐴𝐴𝐴𝐴𝐴𝐴 𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑉𝑅𝑅1 − �(𝑅𝑅1 + 𝑅𝑅2) ∗�5. EXPERIMENTAL RESULTS ⎝�⎞ ⎟⎠(1)To use the circuit described in section 4, it is first required to determine the relationship between temperature and resistance of the microheater. This was accomplished by loading the microheater into a furnace with a thermocouple and recording the resistance of the microheater at various temperatures.

This was accomplished by loading the microheater into a furnace with a thermocouple and recording the resistance of the microheater at various temperatures. Platinum’s resistance changes linearly with changes in temperature [30], so a linear line of best fit was applied to the acquired data. The results of this test are shown in Figure 9.1321120100)Ω(80e c n a t s i s e R6040R(T) = 0.1275T + 44.0122000100200 Temperature (°C)300400Figure 9 - The resistance of the Pt microheater as a function of temperatureFor validation, a PCB of the temperature control circuit was made and the microheater package was connected to it by stacking on top of it. The temperature control board is connected to the Arduino Mega board. This setup can be seen in Figure 10.Figure 10 - Microheater package mounted to control boardOur goal is to bring the microheater to 400°C in 1ms and maintain the temperature using the PID control.

To validate this function, a single microheater was activated and its temperature was calculated using the temperature-resistance relationship shown in Figure 9: the resistance of the microheater was found by measuring the DAC and ADC voltages as shown in Figure 8 and then calculating using Equation 1. This was accomplished with only a maximum of .8W power consumption. The change of temperature and power consumption over time are shown in Figure 11. It shows the designed circuit board can effectively raise and maintain the microheater temperature to the target temperature in the desired timescale.Additionally, because the microheater will be turned on and off to selectively sinter powder particles, it was desired to determine a cycling frequency for the microheater operation for the MAPS process.

For this test, the microheater was heated to the target temperature of 400°C then allowed to cool until reaching room temperature by natural cooling (i.e., no assistive cooling mechanism is used) at which point it is ramped up again, etc. The results of this test are shown in Figure 12. It shows the designed microheater and temperature control circuit can meet the requirements of the MAPS process to quickly turn on and off to reach the target temperature in a couple of milliseconds.13225001)C ° (4000.8)e r u t a r e p m e T3002001000.60.40.2W(r e w o P000123 Time (ms)45TemperatureTarget TemperaturePowerFigure 11 - Pt microheater being heated to a target temperature of 400°C in 1ms500)C ° (400e r u t a r e p m e T3002001000012345Time (ms)TemperatureTarget TemperatureFigure 12 - Cycling the Pt heater on and off resulted in a response time of less than 1ms to heat up and cool down again6.

CONCLUSIONS AND FUTURE WORK In this paper, we have presented the fabrication of a microheater array to be used in the MAPS process. The microheater consists of a platinum filament and gold leads. The microheater was scaled to a 2x2 array and fabricated on a glass substrate, and the die was built with evaporated copper leads on the side for electrical connection and for attachment to the packaging. A feedback controller was devised to control the temperature using a PID control circuit. Tests using this circuit showed that it is possible to bring the heater to 400°C within 1ms and maintain it at that temperature, as originally laid out in the design specification for the MAPS system. In the MAPS system, each microheater can be operated individually to pattern a substrate without the need to move the printhead relative to the substrate; in effect, the area occupied by the microheater array is the printable area of the system.

To accomplish this, a much larger array of microheaters is needed. For future works, a larger array should be fabricated on a glass substrate with through-vias, which will allow for electrical connection on the back side of the substrate.13237. ACKNOWLEDGEMENTSWe gratefully acknowledge the financial support from the Oak Ridge Associated Universitiesthrough the Ralph E. Powe Junior Faculty Enhancement Awards and from the University of Arkansas through the startup fund provided by the Vice Provost Office for Research and Economic Development. Errol Porter and Clint Hardee for their assistance in developing microheater array. Any opinions, findings, and conclusions or recommendations expressed in this publication are those of the authors and do not necessarily reflect the views of the University of Arkansas.REFERENCES[1] A. Van Horn, "Selective Resistive Sintering: A Novel Additive Manufacturing Process," M.S. Thesis,University of Arkansas, Fayetteville, 2016.[2] A.

149, pp. 25-30, 2016.[14] T. Guan and R. Puers, "Thermal Analysis of a Ag/Ti Based Microheater," Procedia Engineering, vol. 5, pp.1356-1359, 2010.[15] A. Pike and J. W. Gardner, "Thermal modelling and characterisation of micropower chemoresistive siliconsensors," Sensors and Actuators B: Chemical, vol. 45, pp. 19-26, 1997.[16] D. Bradley and A. Entwistle, "Determination of the emissivity, for total radiation, of small diameter Platinum- 10% Rhodium wires in the temperature range 600-1450 C," British Journal of Applied Physics, vol. 12, p. 708, 1961.1324[17] S. Joy and J. K. Antony, "Design and simulation of a micro hotplate using COMSOL multiphysics for MEMSbased gas sensor," in Fifth International Conference on Advances in Computing and Communications (ICACC), 2015.[18] G. Velmathi, S. Ramshanker and S.

Bäcklund, "Space qualification of direct bonded siliconmicrosystems," in Micro Nano Technology for Space Applications, Pasadena, CA, 2000.[22] J. Connelly, A. Kourepenis, D. Larsen and T. Marinis, "Inertial MEMS Development for Space," in 2ndinternational conference on integrated Micro/Nanotechnology for Space Applications, 1999.[23] N. Maluf and K. Williams, Introduction to Microelectromechanical Systems Engineering, Artech House,2004.[24] B. Bae, J. Yeom, B. R. Flachsbart and M. A. Shannon, "A Sensorless and Versatile Temperature-ControlSystem for MEMS Microheaters," Transactions of the Korean Institute of Electrical Engineers, vol. 55, no. 11, pp. 544-547, 2006.[25] M. N. H. Z. Alam, A. A. A. Moghadam and A. Kouzani, "Establishment of temperature control scheme for microbioreactor operation using integrated microheater," Microsystem Technologies, vol. 21, no. 2, pp. 415- 428, 2015.[26] J. F. Creemer, D. Briand and H. W.

[Accessed 13 April 2017].[30] "Thermophysical Data on Platinum," Platinum Metals Review, vol. 28, no. 4, p. 164, 1984. [31] HP Development Company, L.P., "Timeline of our history," 2017. [Online]. Available:http://www8.hp.com/us/en/hp-information/about-hp/history/hp-timeline/timeline.html. [Accessed 30 March 2017].[32] M. Baroncini, P. Placidi, G. C. Cardinali and A. Scorzoni, "Thermal characterization of a microheater formicromachined gas sensors," Sensors and Actuators A, vol. 115, no. 1, pp. 8-14, 2004.[33] J. Puigcorbé, D. Vogel and B. Michel, "Thermal and mechanical analysis of micromachined gas sensors,"Journal of Micromechanics and Microengineering, vol. 13, no. 5, pp. 548-556, 2003.1325.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFabrication and Quality Assessment of Thin Fins Built Using Metal Powder Bed Fusion Additive ManufacturingAlexander J. Dunbar¹, Gabrielle J. Gunderman¹, Morgan C. Mader², and Edward W. Reutzel¹¹Applied Research Lab, The Pennsylvania State University (CIMP-3D)²Honeywell | AerospacePowder bed fusion additive manufacturing (PBFAM) is well suited for the fabrication of metallic components with thin features that would be otherwise impractical using other manufacturing methods. As component designs begin to take full advantage of the capabilities of additive manufacturing, so must the capability of measurement techniques used in assessment of quality increase. The characterization of quality may be application specific, requiring different metrics for different uses, e.g. impact on thermal vs. mechanical considerations.

The characterization of quality may be application specific, requiring different metrics for different uses, e.g. impact on thermal vs. mechanical considerations. Here, thin fins are built with a wide range of processing conditions to test the limits of thin, metallic components using PBFAM. These thin fins are inspected using novel computed tomography (CT) based measurement techniques to assess their build quality. Within the process parameters tested, fins were successfully built thinner than manufacturer-recommended minimum wall thickness using default process parameters. The quality of these fins was assessed utilizing post-build non-destructive evaluation techniques developed herein.Keywords: Powder bed fusion, CT, Inconel 625, Heat exchangerIntroductionPowder bed fusion additive manufacturing (PBFAM) technology enables unique designs that revolutionize many of the traditional manufacturing design guidelines and principles.

In particular, PBFAM provides the opportunity to design and fabricate structures with thin internal structures that would otherwise be impossible for more traditional subtractive manufacturing methods. Of particular interest is the use of PBFAM in the design and construction of heat exchangers, which benefit from light weighting and advanced geometries [1], [2].Additive manufacturing is well suited for the manufacture of heat transfer components. Thoughfor many applications the surface roughness present of as-built PBFAM components represents a hindrance in its application due to reduction in fatigue properties [3], for heat transfer applications, the increased surface roughness can be of benefit in enhancing convective heat transfer [4]–[6].

Potential improvements to heat exchangers with PBFAM technology are not limited to increased surface roughness, but also include the complex and thin geometries made viable through its use [5], [7], [8].While PBFAM is well suited to the design and fabrication of thin geometries that would be found within a heat exchanger, thin geometries may result in further complications. While commercial PBFAM systems come with material specific processing parameters for the fabrication of components, previous studies have shown that these parameters may not be ubiquitous to all part geometries [9], [10]. Yadroitsev et al. (2007) demonstrated that effective processing parameters may also be dependent not only on material, but also local geometry [11].

Yadroitsev et al. (2007) demonstrated that effective processing parameters may also be dependent not only on material, but also local geometry [11]. Based on these studies it is imperative to define process parameters that are specific to not only machine and material, but geometry as well.Prior studies have shown that PBFAM heat exchangers can exceed the heat transfer capabilities of traditionally manufactured ones [8], [12], [13]. Further advancements are possible through redesign of heat exchangers specifically for additive manufacturing [13]. As part of this effort it is necessary to identify PBFAM system capabilities and define processing parameters that will produce optimal heat2650exchanger features. This study seeks to define the limits of a PBFAM system in producing the thinnest fins that would be suitable for use in a heat exchanger.

This study seeks to define the limits of a PBFAM system in producing the thinnest fins that would be suitable for use in a heat exchanger. Fins are measured using novel X-Ray computed tomography (CT) based measurement techniques developed herein; measurements are compared to traditional cross-sectioning measurements.MethodologyThin fins at a range of different fin angles were all constructed on an EOS M280 PBFAM machine using Inconel 625 with 40 µm layers. Fin geometries, shown in Figure 1, include a pedestal to isolate the fin geometry from the build plate and to allow wire EDM removal of the fins without damage. The top surface of the pedestal is angled so that the fin is always built perpendicular to the surface. The fin angles, defined as the angle at which the fin is built with respect to the (X-Y) build plane, are varied (45°, 67.5°, and 90°) to assess the impact of fin angle on quality.

The fin angles, defined as the angle at which the fin is built with respect to the (X-Y) build plane, are varied (45°, 67.5°, and 90°) to assess the impact of fin angle on quality. Fins are built using a scan path, with single- double and offset-scan trajectories, along the length of the fin using a range of laser powers and scan velocities. To achieve this the fins themselves are designed as a zero thickness 2D surface geometry, and ascribing user-specified scan parameters to each fin.Figure 1: Schematic illustrating the fin and pedestal geometry and dimensions (in mm). All fins extend 9.65 mm into the page. (A) Fin angle 45° (B) Fin angle 67.5° (C) Fin angle 90°2651Figure 2: Schematic defining build angle for fins with the build direction perpendicular to the XY planeRecognizing that recoater blade interference can also play a role in the quality of thin features, thebuild angle, defined as the angle between the fin and the recoater, as shown in Figure 2, was also varied.

Only two build angles were used, where a build angle of 90° is defined as being perpendicular to the recoater and a build angle of 45° is defined as the fin being rotated 45° about the Z axis. Fins were spaced evenly with a constant radius about the center of the build plate, as shown in Figure 3 for the 45° build. A constant radius was used to reduce effects that may arise from variations in laser path length and angle-of-incidence. A summary of geometric parameters that are varied in this study can be found in Table 1.Figure 3: Build layout for fins built with 45° build angleTable 1: Summary of geometric parameters varied in the fin studyBuild Angle (°)Fin Angle (°)45459067.5902652Processing parameters altered within this study include: scan speed, laser power and the laserscan type. Power and scan speed are common parameters to vary in the definition of new process parameters [11].

Power and scan speed are common parameters to vary in the definition of new process parameters [11]. The laser scan type determines the number of scan paths for each layer and the relative distances between the scan paths. For these fins, the laser scan type is used in place of the more common parameter, hatch spacing. Three different laser scan types are used: single, double, and offset. Fins built with the single laser scan type are built with only one laser path per layer. Fins built with the double laser scan type are built with two laser scan paths with each overlapped directly on one another. The final laser scan type, referred to as offset, is defined by two laser trajectories that are offset by 70 µm.

The final laser scan type, referred to as offset, is defined by two laser trajectories that are offset by 70 µm. The various process parameters used in the study are summarized in Table 2.Table 2: Summary of process parameters used in fin studyScan Speed (mm/s)Laser Power (W)Laser Scan Type90070Single (S)960100Double (D)1500285Offset (O)Due to the large number of processing and geometric parameters within this study, a completedesign of experiment was not feasible. Instead, fin angle and laser scan type parameters were varied together to enable assessment across a larger parameter matrix. Table 3 shows a list of the parameter sets used for each build, i.e.

Each of the default fins were designed to be a specified thickness: 300 µm, 200 µm, 100 µm, and 75 µm. Laser tool paths were defined automatically with fin geometries receiving the default laser scan path trajectories as defined by the EOS process software (PSW). These fins were built at two different build angles (45° and 90°) and only one fin angle (90°).All fins from the study are subjected to 3D computed tomographic (CT) scanning using a GEphoenix v|tome|x system with a 300 kV microfocus tube. CT scan parameters were selected to provide 15 µm voxels for the fins. The CT scan data for each fin was processed using a thresholding technique to differentiate solid material. The threshold value was determined through a qualitative analysis. Fortunately, as the fins are very thin, scans have a high contrast between solid and void regions which provides a clear definition of both solid material.

Fortunately, as the fins are very thin, scans have a high contrast between solid and void regions which provides a clear definition of both solid material. An example cross-section, comparing greyscale and binary images, for a fin is shown in Figure 4.Figure 4: Comparison of greyscale (left) and binary (right) for a single fin slice.Using the thresholded data, the median thickness and through-hole porosity of each fin werecalculated within a user-defined central region of interest within each fin (to eliminate edge effects). The region of interest is 5mm wide and 2 mm tall and is centered on the fin. After CT scans were completed, each fin was also cross-sectioned, mounted, and polished in a destructive process for comparison. The thickness of each cross-sectioned fin was measured at five locations along the height. An example fin cross-sectioning measurement is shown in Figure 5 for fin H, which has a 67.5° fin angle and a 45° build angle.

An example fin cross-sectioning measurement is shown in Figure 5 for fin H, which has a 67.5° fin angle and a 45° build angle. Although cross-sectional imaging is a more conventional measurement technique, when trying to assess the true thickness of the fin, it can be limiting. CT-based fin quality metrics allow for thickness and through-hole porosity measurements to be aggregated across a larger region of the fin area instead of confining measurements to individual cross-sectional areas.2654Figure 5: Fin H (Build angle 45°, Scan speed 900 mm/s, Power 285 W, Fin angle 67.5° and Laser scan type Single) as built image (left) and cross-sectional measurements (right) (Measurements in Imperial Units)2 mm5 mmFigure 6: Single CT image of fin H with measurement regions highlighted.

Region of interest (ROI) shown as a red rectangle and CT-based cross section region shown as a green line.For thin AM components there is a great deal of variability in both through-hole porosity andthickness. For comparison, the CT-based measurements are taken for fin H across a single cross-section, defined by the green line in Figure 6, and a region of interest (ROI), defined by the red rectangle in Figure 6. CT-based cross sectioning and CT-based ROI measurements are compared against manual cross- sectioning measurements in Figure 7. Table 4 shows the comparisons of median thickness of the three different measurement techniques used. Between the three measurement techniques there is a significant discrepancy in the median thickness. Even between the two CT-based measurements, Figure 7-(B) and Figure 7-(C), there is a discrepancy of 30 µm in the median thickness.

Even between the two CT-based measurements, Figure 7-(B) and Figure 7-(C), there is a discrepancy of 30 µm in the median thickness. Figure 8, a contour map of fin thickness across the ROI, shows the variation in thickness across the region of interest. This discrepancy is not the result of measurement error, but instead a result of the complex surface in the relative2655measurement region.

The high variability of thickness across the fin defines the need for metrics that are aggregated across the surface of the fin to better asses the build quality.Figure 7: Histograms of measured fin thickness using different methods: (A) Manual Cross-Sectioning (Destructive) (B) CT- based measurement aggregated across a single cross-section (C) CT-based aggregated across an entire ROI.Table 4: Comparison of measurement techniques for median thickness of a fin.Measurement TechniquePhysical Cross-Sectioning CT Cross-Sectioning CT Region of InterestNumber of Measurements Per Fin 5-6 ~500 ~40,000Median Thickness (µm) 267.5 190 2252656Figure 8: Contour map of thickness across the region of interest for fin H (as measured from CT scan data).ResultsA comparison of the median thickness measurements from the CT based analysis and the cross-sectioning measurement method is shown in Figure 9.

The comparison between the manual cross- sectioning measurements and CT-based ROI measurements reveal that the CT-based measurements frequently exceed measurements from manual cross-sectioning. Although the percent difference in some of these cases exceeds 30% it is important to note that discrepancies can be due to both differences in measurement techniques as well as what is also likely due to differences in the region of the fin where the measurements are taken.

The consistent agreement between the two methods provides adequate confidence in the CT-based ROI method, therefore all further analysis will be respective to CT measurements.Figure 9: Comparison of median fin thickness from CT-based ROI measurement (CT) with manual cross-sectioning measurement (Cross-Sect) for build angle 45° (left) and 90° (right)2657Figure 10:Percent difference in the median fin thickness between CT-based ROI measurements and manual cross-sectioning for build angle 45° (left) and 90° (right)Figure 11 and Figure 12 show the median thickness and through-hole porosity, respectively, forall fins built within this study. Fins that failed to build are designated in Figure 12 as having 100% through-hole porosity. Fin thicknesses reported in Figure 11 demonstrate that for a given scan speed (V) there is an increase in fin thickness as laser power increases.

Fin thicknesses reported in Figure 11 demonstrate that for a given scan speed (V) there is an increase in fin thickness as laser power increases. Additionally, there is a second trend that shows that as the scan speed decreases the resultant fin thickness increases. Neither of these trends are surprising as an increase in laser power or decrease in scan speed result in more energy input to a given region. Increased energy deposited into a given volume will increase the melt pool size and therefore produce a thicker fin. Fins built with a build angle of 90° are consistently thicker than the matching fin built with a build angle of 45°. Figure 12, in conjunction with Figure 11, show that there is a trend where the thinner fins also exhibit the highest levels of through-hole porosity.

Figure 12, in conjunction with Figure 11, show that there is a trend where the thinner fins also exhibit the highest levels of through-hole porosity. The through-hole porosity and median thickness are also reported in Table 6.The thinnest fins produced with no porosity were either scanned with a single or a double laserscan, since the overlap scan strategy consistently produced thicker fins. This is consistent with the initial supposition due to the increased area being scanned by the laser due to offset of 70 µm for that laser scan type. Sorting fins by highest through-hole porosity identifies that the single laser scan is most likely to exhibit a higher porosity, followed by the double laser scan, with the overlap laser scan having the lowest porosity.

It should be noted that the single, double, and overlap laser scan types were not isolated parameters in the test plan defined in Table 3.2658Figure 11: Median fin thickness for all fins built highlighting the laser power (P) and scan speed (V) parameters for each fin. For the values corresponding to high medium and low see Table 3.Figure 12: Percent through-hole porosity for all fins built highlighting the laser power (P) and scan speed (V) parameters for each fin. For the values corresponding to high medium and low see Table 3.Figure 13 shows the median thickness and through-hole porosity of the fins built with the default EOS processing parameters and prescribed fin thickness defined in the Methodology section. In Figure 13 there is a discrepancy between the prescribed fin thickness and the measured thickness.

In Figure 13 there is a discrepancy between the prescribed fin thickness and the measured thickness. For the fins with 300 µm prescribed thickness, the measured thickness is approximately 300 µm, but the fins designed with 200 µm, 100 µm and 75 µm wall thickness exhibit similar thicknesses and through-hole porosity. This is likely an effect of the scan strategy, where only the 300 µm thick fin has a hatching region and the discrepancy between the other fins is defined by the offset of the contours tracing the outline of the fin as defined by the EOS process software. Typical hatch spacing for a PBF machine is between 100 µm and 200 µm, which here would allow for hatches in the 300 µm fins.2659Figure 13: Median thickness and through-hole porosity of fins built with default EOS processing parameters and prescribed thicknessDiscussion For fins within the channel section of a heat exchanger (i.e.

with the same fluid on either side of the fin), some porosity is acceptable as flow across the fin surface would not result in unwanted mixing. However, for a tube sheet, which separates two different fluids, the geometry must be capable of preventing mixing and sustaining a pressure differential between the two flows; which requires that no porosity is present. In this analysis, we have addressed this by defining two categories of fins: the thinnest fins with less than 10% porosity, and the thinnest fins with no porosity. Figure 14 highlights fins that match these definitions of successful fins.Figure 14 shows median fin thickness for all fins that are built with no through-hole porosity (left) and meet the 10% through-hole porosity threshold (right). For the fins built with the parameters defined in Table 3, the thinnest fin with no porosity has a median thickness of 97.5 µm and the thinnest fin within the 10% porosity threshold has a median thickness of 75 µm.

For the best fins there is also a strong agreement, within 10% difference, between the cross-sectioning measurements and the CT-based ROI measurements. For fins built with default EOS parameters, Figure 13, the only fins built with no porosity has a median thickness of 338 µm. The thinnest fin that meets the 10% porosity threshold has a median thickness of 113 µm.

Within the parameter space tested, the fin thickness increases with either an increase in laser power or a decrease of scan speed. Fins were analyzed with a combination of CT-based measurements aggregated over a region of interest, developed herein, as well as physical measurements performed using destructive cross-sectioning. The CT-based measurement technique represents an advancement in inspection methods allowing geometric measurements to be made automatically, allowing for a larger and more representative region of interest to be used to quantify build quality.Fins that were built with the parameters defined in Table 3 were compared against fins built with default EOS processing conditions. The best fins built with the varied processing conditions were thinner and less porous than the fins built with the default EOS processing conditions.

The best fins built with the varied processing conditions were thinner and less porous than the fins built with the default EOS processing conditions. Further development includes using the identified ideal processing parameters for fins in actual heat exchangers to test their effectiveness in working components. Future work also includes the development of additional measurement algorithms to determine other pertinent metrics to define fin quality using CT-based measurement techniques.AcknowledgementsThe authors gratefully acknowledge the contributions of Griffin Jones and Corey J. Dickman. This material is based on research sponsored by Air Force Research Laboratory under agreement number2661FA8650-12-2-7230. The U.S. Government is authorized to reproduce and distribute reprints for Governmental purposes notwithstanding any copyright notation thereon.

brication and Testing of Graphite Bipolar Plates for Direct Methanol Fuel CellsKaushik Alayavalli and David L. Bourell,Laboratory for Freeform Fabrication, The University of Texas at Austin,Austin, Texas 78712AbstractBipolar plates made from graphite are highly desirable in fuel cell applications due to the properties of high electrical conductivity, low weight and resistance to corrosion that graphite possesses. However, the poor mechanical properties of natural graphite lead to prohibitive machining costs. Graphite parts are made by indirect SLS using graphite powder mixed with a phenolic resin binder which is subsequently burnt out. These parts fabricated using indirect SLS have poor green strength and electrical conductivity. The current work deals with the addition of carbon fiber to the graphite-phenolic mixture with an aim to improve green part strength and brown part conductivity.

The current work deals with the addition of carbon fiber to the graphite-phenolic mixture with an aim to improve green part strength and brown part conductivity. This research work also addresses fabrication of graphite bipolar plates for Direct Methanol Fuel Cells (DMFCs) with internal features and flow channels with different cross-sectional geometry that are difficult to produce using conventional machining processes.1. IntroductionFuel cells are a clean energy technology with the benefits of high efficiency energy conversion and low power source emission. Most attractive is the high efficiencies achievable at temperatures under 100 ºC [1]. The basic operating principle of any fuel cell is the conversion of the chemical energy of the supplied fuel into electrical energy. Direct Methanol Fuel Cells (DMFCs) are a type of Polymer Electrolyte Fuel Cells (PEMFCs) that utilize methanol as a fuel.

Direct Methanol Fuel Cells (DMFCs) are a type of Polymer Electrolyte Fuel Cells (PEMFCs) that utilize methanol as a fuel. The energy density of methanol which is higher than that of Li – ion batteries coupled with the fact that methanol is a readily available liquid which is easy to store and transport, makes it an attractive option for fuel cell applications. However, DMFC technology is hampered by issues such as methanol crossover [2] from the anode to the cathode side of the fuel cell and unstable electrocatalysts [3].In a DMFC the supplied fuel is a mixture of methanol and water which is oxidized into CO2 and H+ ions. The H+ ions are transported across a polymer electrolyte membrane to the cathode side where they react with the supplied oxygen to form water. Figure 1 shows the schematic of a DMFC single cell, which comprises of membrane electrode assembly (MEA), sandwiched between bipolar plates that supply the fuel and provide structural support for the fuel cell stack.

In a stack, bipolar plates serve the purposes of fuel distribution, current collection, separation ofthe membrane electrode assemblies and bearing of the clamping force holding the fuel cell stack together. They are the most significant components of a fuel cell stack in terms of weight and cost of production [4]. Graphite has been the standard material of choice for bipolar plates due to its high electrical conductivity and low mass. However, machining fine features such as flow channels and fuel manifolds on graphite plates is expensive since graphite is brittle and has poor mechanical properties [5].The performance of the fuel cell is dependent on factors such as the efficiency of the catalysis process, the electrical properties of the cell and the flow field design. Due to the difficulties in machining, testing of novel and complex flow field designs is an expensive, time consuming process.

Due to the difficulties in machining, testing of novel and complex flow field designs is an expensive, time consuming process. Indirect selective laser sintering (SLS) offers the benefit of producing graphite plates with features of significant complexity in a single step. Indirect SLS of graphite bipolar plates involves mixing graphite powder with a binder such as finely milled phenol formaldehyde (“phenolic”) powder and using a laser to scan individual layers of a .stl file into a solid 3-dimensional object. The ‘green’ SLS parts produced are porous and require post processing for the improvement of electrical properties and complete densification.

The ‘green’ SLS parts produced are porous and require post processing for the improvement of electrical properties and complete densification. The post processing steps include heating the green parts in a vacuum furnace to burn off the phenolic binder and infiltrating the porous ‘brown’ part with a suitable infiltrant to densify and strengthen the part.Due to the presence of fine features in the flow field of the bipolar plate, there is a need for the green parts to be able to withstand the break out process without being damaged. The current work, investigates the suitability of carbon fiber as an additive to the graphite-phenolic mixture to improve green strength and overall conductivity. Three flow field designs are also evaluated for performance and verification of SLS as a process that can produce repeatable fuel cell results.2.

Three flow field designs are also evaluated for performance and verification of SLS as a process that can produce repeatable fuel cell results.2. Experimental Process2.1 Bipolar plate production by Indirect Selective Laser SinteringThe process of producing graphite bipolar plates by indirect SLS is shown in figure 2. It involves the production of a green part by the indirect SLS of a mixture of graphite (~ 70 microns, GS-150E, Graftech Inc) and phenolic (~ 11 microns, GP-5546, Georgia Pacific) powders in a 70 wt% to 30 wt% ratio, in a Sinterstation 2000 [6]. The porous green part is then heated in a vacuum furnace at temperatures greater than 1200 ºC to carbonize the phenolic binder, thus producing a ‘brown’ part. Phenolic is a “high ash” polymer, producing significant residual carbon after thermal dissociation. This carbon is effectively a high-temperature binder of the graphite powder.

Phenolic is a “high ash” polymer, producing significant residual carbon after thermal dissociation. This carbon is effectively a high-temperature binder of the graphite powder. The parameters used for the SLS process are: fill laser power (10 W – 20 W), outline laser power (4 W), powder layer thickness (0.1016 mm), laser scan spacing (0.0762 mm), laser scan speed (1.524 m/s). Based on Differential Scanning Calorimetry (DSC) results of the phenolic powder, the part bed and feed bins were maintained at 65 ºC and 47 ºC, respectively, to ensure that the phenolic powder bed that was not laser sintered remained stable without being thermally affected.

Post processing of this brown part includes infiltration with low viscosity cyanoacrylate glue [7] that seals the pores, thereby making the part fluid impermeable, and polishing of the surfaces to expose the underlying electrically conductive graphite matrix.Powder Preparation & MixingIndirect SLS of Green PartVacuum Furnace CarbonizationLow viscosity polymer infiltrationRoom temperature cure/ PolishingMaterial Properties Characterization and TestingFinished Bipolar PlateFigure 2. Flow chart of the experimental process of bipolar plate production by indirect SLS2.2 Carbon fiber additiveCarbon fiber (length ~ 70 – 100 µm, diameter ~ 7 µm) was added to the graphite – phenolic resin mixture in three compositions, 10%, 20% and 40% by weight w.r.t the weight of graphite powder. The amount of phenolic binder was kept constant at 30% by weight.

The amount of phenolic binder was kept constant at 30% by weight. Three point bend specimens (127 mm x 13 mm x 3 mm) were built in an x-y orientation using the same SLS parameters as the regular graphite – phenolic mixture. The electrical conductivity of these specimens was first measured (ASTM C611) using a Cascade Microtech four point probe station in the green stage and the flexural strength of the specimens was then obtained using a threepoint bend method (ASTM D790). Green parts were also burnt out in batches of three at temperatures varying from 400 ºC to 1200 ºC to produce brown parts whose electrical conductivity and flexural strengths were then measured.Figure 3.

SEM images of the carbon fiber at a) 1000x and b) 4500x magnification.2.3 Flow field designs for improved fuel cell performanceThe design of the bipolar plate flow field has a significant impact on the various factors that affect the performance of the DMFC such as pressure drop across the cell, fuel absorption in the catalyst layers and transport of CO2 and water out of the cell to name a few.

Studies in flow field design [8, 9, 10] have shown that the cross-sectional geometry of the channel, aspect ratio of the channels (ratio of channel width to rib width) and the flow field pattern (serpentine, corrugated, interdigitated etc) all affect the performance of the fuel cell.The electrochemical performances of the DMFC was evaluated with a single-cell fixture having active areas of 5 cm2 and 10 cm2and feeding a preheated methanol solution into the anode at a flow rate of 2.5 mL/min by a peristaltic pump without back pressurization and humidified oxygen into the cathode at a flow rate of 200 mL/min at ambient pressure [11]. Three basic shapes, elliptical, triangular and rectangular (figure 4) were chosen as the cross-sectional geometries of the flow field channel for comparative purposes. The flow field area was chosen to be 5 cm2 and two sets of the geometries were made differing in the number of passes (10 and 15) of the flow field.

The flow field area was chosen to be 5 cm2 and two sets of the geometries were made differing in the number of passes (10 and 15) of the flow field. One additional set of 10 cm2 flow field area (10 passes) was made to verify the results on scaling up. These plates were post processed and tested in a Scribner Associates Inc. 850e fuel cell test station with the following commercially purchased MEA: Electrolyte Membrane: NaFion 117, Anode Catalyst: Pt/Ru/C, Cathode Catalyst: Pt/C, (4 mg/cm2 loading). The performance of the plates made by SLS was compared to the performance of commercially purchased synthetic graphite plates with a rectangular cross-section and a serpentine flow field of 5 cm2 making 15 passes.(a)(d)(b)(e)(c)(f)Figure 4 a) elliptical, b) triangular and c) rectangular flow fields with 10 passes.Figure 4 d) elliptical, e) triangular and f) rectangular flow fields with 15 passes.3.

Results and Discussion3.1 Flexural strength of carbon fiber added partsAs seen from figure 5a, the flexural strength of the green parts increased with an increase in the percentage of carbon fiber added. An improvement in the strength of the green parts was not apparent with a 10% addition of carbon fiber. The difference in flexural strength values is small enough to be within the margin of error of the test process. Adding 20% to 40% of carbon fiber shows a significant increase (50% and 100% respectively) in the flexural strength of the green parts.Figure 5 a) Flexural strength of green parts w.r.t addition of carbon fiber.Figure 5 b) Flexural strength of brown parts w.r.t burnout temperature.The flexural strength of the brown parts was found to dependent on the burnout temperature (figure 5 b). As in the green parts, the strength of the brown parts was found to increase with an increase in the amount of carbon fiber added.

As in the green parts, the strength of the brown parts was found to increase with an increase in the amount of carbon fiber added. The flexural strength of the parts burnt out at 600 ºC was found to consistently lower than those sintered at 400 ºC. This could be due to incomplete dissociation of the phenolic binder. From 600 ºC to 1200 ºC, the increase in flexural strength is seen to be almost linear. There is almost a 90% increase in the strength of the brownparts with 40% carbon fiber (9.933 MPa) burnt out at 1200 ºC compared to the brown parts without any carbon fiber ( ~ 5.27 MPa) burnt out at the same temperature.3.2 Electrical conductivity of carbon fiber added partsThe electrical conductivity of the brown parts were found to increase with sintering temperature for the samples with 10% and 20% carbon fiber added (figure 6). For the samples with 40% carbon fiber, the conductivity increased as expected until 800 ºC and then decreased at temperatures above 1000 ºC.

For the samples with 40% carbon fiber, the conductivity increased as expected until 800 ºC and then decreased at temperatures above 1000 ºC. In general, compared to samples without the any carbon fiber, the electrical conductivity of carbon fiber added graphite parts is lower. This may be due to a lack of interconnectivity in the longitudinal fiber axis due to the small size of the fibers and also the percolation threshold.

This may be due to a lack of interconnectivity in the longitudinal fiber axis due to the small size of the fibers and also the percolation threshold. It has been shown that the electrical conductivity of carbon fiber filled composites is strongly dependent on the orientation of the fibers as well the percolation threshold [13, 14].Figure 6 a) Electrical conductivity of graphite- phenolic parts w.r.t burnout temperature.Figure 6 b) Electrical conductivity of carbon fiber added parts w.r.t burnout temperature.3.3 Fuel cell performance testsThe performance of the 10 pass single serpentine plates with flow field areas of 5 cm2 and 10 cm2 was seen in general to be comparable to that of the commercially procured 15 pass single serpentine 5 cm2 plates (figure 7a, 7b and 7c). Due to the superior electrical properties of synthetic graphite, the commercial plates showed better performance in the ohmic region of the polarization curve.

Due to the superior electrical properties of synthetic graphite, the commercial plates showed better performance in the ohmic region of the polarization curve. The 5 cm2 SLS plates with triangular and elliptical cross-sections, performed comparable to the commercial plates while the plates with a rectangular cross-section showed the poorest performance in both the ohmic and mass transport regions (figure 7a). This performancewas also seen with the 10 cm2 plates, where the triangular and elliptical cross-sections outperformed the plates with the rectangular cross-section (figure 7b).The performance of the SLS plates with 15 passes (figure 7c) was comparable to the commercial plates in the case of the rectangular cross-section. The rectangular cross-section plates showed similar performance against the triangular and elliptical cross-section as in the case of the plates with 10 passes.

The rectangular cross-section plates showed similar performance against the triangular and elliptical cross-section as in the case of the plates with 10 passes. The plates with triangular and elliptical cross-sections outperformed the commercial plates in the mass transport region (higher current densities) while performing almost the same in the ohmic region. This suggests that the elliptical and triangular cross-sections enable the methanol to reach catalyst sites more readily than the rectangular cross- sectioned SLS and commercial plates.Figure 7 a) DMFC polarization curves of 5 cm2 flow field area bipolar plates (10 passes)Figure 7 b) DMFC polarization curves of 10 cm2 flow field area bipolar plates (10 passes)Figure 7 c) DMFC polarization curves of 5 cm2 flow field area bipolar plates (15 passes)4. ConclusionsBipolar plates produced by indirect SLS were shown to perform comparable to and in some cases better than commercially produced synthetic graphite plates.

ConclusionsBipolar plates produced by indirect SLS were shown to perform comparable to and in some cases better than commercially produced synthetic graphite plates. It was found that flow fields with triangular and elliptical cross-sections showed better mass transport performance than flow fields with rectangular cross-sections. Carbon fiber was used as an additive in the SLS process and found to improve both the green part and brown part strengths. A 100% improvement in green part strength was obtained, while a part with 40% carbon fiber, sintered at 1200 ºC, was found to have almost 90% more flexural strength than parts made without carbon fiber. However, it was observed that adding carbon fiber reduces the electrical conductivities of the brown parts compared to parts made with just graphite and phenolic.

However, it was observed that adding carbon fiber reduces the electrical conductivities of the brown parts compared to parts made with just graphite and phenolic. This poses a tradeoff, in terms of optimum green part strength and brown part conductivity, which needs to be evaluated depending on the bipolar plate design.AcknowledgementThis work was supported by the Office of Naval Research MURI grant No. N00014-07-1-0758.REFERENCES[1] Vielstich, W., Lamm, A., and Gasteiger, H. (2003), Handbook of Fuel Cells, Vol. 1, Wiley and Sons Ltd., London.[2] Casalegno, A., Marchesi, R. (2008), DMFC performance and methanol cross-over: Experimental analysis and model validation, Journal of Power Sources, Vol. 185, pp.318–330[3] Liu, H., Songa, C., Zhang, L., Zhang, J., Wang, H., Wilkinson, D. P. (2006), A review of anode catalysis in the direct methanol fuel cell, Journal of Power Sources, Vol. 155, pp.95–110[4] Mehta, V. and Cooper, J.S.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFabrication of 3D Multi-material Parts Using Laser-based Powder Bed FusionC. Anstaett*, Dr.-Ing. C. Seidel*, Prof. Dr.-Ing G. Reinhart* * Fraunhofer Research Institution for Casting, Composite and Processing Technology IGCV, Augsburg, GermanyAbstractAs different branches of industry use Laser Beam Melting (LBM) and more and more materials can be produced with it, this technology goes in and out of focus of production technologies in an industrial environment. A big advantage of LBM is the possibility of building very complex parts and therefore minimizing the need for raw material.

A big advantage of LBM is the possibility of building very complex parts and therefore minimizing the need for raw material. The effects of this, such as lightweight-design, resource-efficient production and reduction of manufacturing time, can even be increased if material can be used locally defined in a part, so that one part does not have to consist of one material per fabrication-process, but at least two. Since LBM is a powder-bed-based process, the implementation of this idea is possible by adopting the conventional delivery device and the manufacturing process itself. In this paper the results of a multi-material process are shown and the influences of different material properties on the manufacturing process are derived.IntroductionIn recent years multi-material has been in the scope of different research groups all over the world, especially in case of additive manufacturing.

Thus far there are already ways to build multi-material parts out of plastic, using this technology for example, by using fused layer modeling (FDM) or stereolithography [1]. Also, in the case of parts made of metal, there are many approaches using laser powder depositioning, for example [2]. Even for powder-bed-based technologies there are many investigations being done concerning the powder depositioning [3–5, 5–7]. However, there is a big difference in the understanding of multi-material, especially in the case of laser beam based powder-bed fusion (LBM). They all have in common the combining of different materials, but vary in the way two or more materials are joined during the manufacturing process.

They all have in common the combining of different materials, but vary in the way two or more materials are joined during the manufacturing process. In order to understand multi-material, Figure 1 shows the different kinds, as well as a mono-material part to more clearly point out the differences.Mono-material2D Hybrid2D Multi-material3D Multi-materialn o i t c e r i dg n dil iu BSubstrateSubstrateSubstrateSubstrateFigure 1: Parts categorized by the meaning of multi-materialA 2D Hybrid part differs from a mono-material part in the material of the substrate, which is the basis for the build-up process. While a mono-material process uses the same material for powder material as for the substrate, a 2D Hybrid process builds on a different material. It is also possible to use a conventionally manufactured part as substrate and start the additive building processes there. Hence, it is called a hybrid part, since different manufacturing techniques are used [8].

Hence, it is called a hybrid part, since different manufacturing techniques are used [8]. One example is shown by Fu et al. who built a copper- based alloy on a steel substrate [9]. One step further is to use LBM for production of both parts and build one material on top of the other. So a 2D Multi-material part can be produced, like Sing et al. has done [10]. This is called a 2D Multi-material part since the change of material occurs only in one dimension perpendicular to the building platform in direction of building. The highest level of multi-material parts consist of multiple materials distributed arbitrarily in the part in all dimensions – a 3D Multi-material part.1548No matter what type of multi-material part is going to be built, the building process is influenced, since all materials vary in their material properties. Therefore the process parameters need to be adapted.

Therefore the process parameters need to be adapted. A 3D Multi- material part has even more influence, since the process chain has to be modified in order to allocate and solidify different materials. In the next sections a methodical approach to build a 3D Multi-material part is shown using a copper-based alloy and tool steel 1.2709.Experimental SetupThere are different approaches for integrating the second powder within the process in order to build a 3D Multi-material part [11]. An approach based on the conventional delivery device was used for this work, supplemented by a suction-module [3, 12, 13].

An approach based on the conventional delivery device was used for this work, supplemented by a suction-module [3, 12, 13]. The process steps is shown in Figure 2.Low ering Building Plat f orm6.1.Pow der Delivery M at erial ASolidif icat ion M at erial B5.2.Solidif icat ion M at erial APow der Delivery M at erial B4.3.Removing unsolidif ied M at erial AFigure 2: Process steps of the multi-material-LBM ProcessAfter delivering and solidifying the material A, the non-solidified particles get removed via a suction- module. Afterwards the second powder of material B is delivered and solidified. This powder remains in the building chamber for the rest of the building process.For this study an SLM 250HL machine from SLM Solutions was used. It is equipped with a fiber laser with a maximal power of 400 W and focal diameter of 100 μm.

It is equipped with a fiber laser with a maximal power of 400 W and focal diameter of 100 μm. In addition, the following gas-atomized powder materials were utilized: Tool steel 1.2709 (X3NiCoMoTi18-9-5) [14] with a normal distribution between 10 and 45 μm, and 2.1293 CuCr1Zr (in the following called CCZ) [15], with a normal distribution between 20 and 45 μm. Both materials have a spherical shape. The chemical contribution of each material is shown in Table 1 and Table 2. Every sample was built with a layer thickness of 30 μm. The specimens were examined under a Nikon SMZ 1000 light microscope and a Keyence VK-9710 laser microscope. The density of the mono-material was tested using two different methods. First, the Archimedes method was used, which considers the entire specimen.

The density of the mono-material was tested using two different methods. First, the Archimedes method was used, which considers the entire specimen. Additionally, the porosity of polished samples was examined by microscopic analysis to get a comparability for the multi-material parts, which cannot be tested by the Archimedes method due to the unknown ratio of different materials [16, 17].

The material distribution within a 3D Multi-material part was measured with an EDX-System of Brucker Quantax 70 via a Hitatchi TM 3030 Plus scanning electron microscope.Table 1 Chemical composition of 1.2709C ≤0.03Mo 5.0Ni 183.0Co 10.0Ti 1.0Table 2 Chemical composition of CCZCr 0.5-1.2Zr 0.03-0.3Fe ≤0.08Si 0.1other constituents ≤0.21549Fe restCu restMethodical ApproachIn order to produce parts with a high density and to derive relationships and influences of the process, amethodical approach represented in Figure 3 was used.Mono-material single-origin materialCCZ1.2709CCZ1.2709CCZ1.2709CCZ1.27092D Multi-material with discrete material transitionMono-material with mixed material25-7550-5075-253D Multi- material1.2709zCCZxFigure 3: Methodical approach for building 3D Multi-material partsFirst of all, the mono-materials are qualified for the LBM process.

Based on existing publications for1.2709 [7, 18] and CCZ [9, 19, 20], the proceeding promoted by [21] was used. However, as information already existed, building of single lines and areas was skipped and only the hatch spacing, scan speed and laser power were varied in the field of the presented parameters.In a second step, parts with two 2D material transitions, which look like sandwich structures, are built to investigate the influence of different material properties on the transition area, e.g. the metallurgical appearance. At the same time, parts with a mixture of both materials are produced to analyze the resulting material distribution and defining parameters for a transition area when building 3D multi-material parts. Here laser power and scanning speed are varied based on the evaluated mono-material-parameters, so that the resulting energy corresponds to the material distribution.

Here laser power and scanning speed are varied based on the evaluated mono-material-parameters, so that the resulting energy corresponds to the material distribution. There were three different mixing ratios under research, shown in Table 3.Table 3 Mixing composition for building the mono-material parts with mixed powderCCZ 25% 50% 75%1.2709 75% 50% 25%These parts will also be analyzed by EDX and later compared to results of the 3D Multi-material parts inorder to analyze the composition of the overlap area.Finally, the 3D Multi-material parts are built. Based on the knowledge gained from previous investigations and a pretest published at EuroPM 2017 [12], 3D Multi-material parts were built using the found parameters. The pretests have already shown that it is necessary to have an overlap between both material areas in order to build a connecting part [12]. Thus the following parts were all built using the found parameter sets and have an overlap of 0.2 mm (Figure 4, left).

Thus the following parts were all built using the found parameter sets and have an overlap of 0.2 mm (Figure 4, left). The focus of investigation is on the building order, the material solidification order and the cleanliness of each material area. In this case the building order means that the inner material – the core – or the surrounding material – the shell – is solidified first (Figure 4, right). The solidification order represents if 1.2709 or CCZ is first solidified.

The solidification order represents if 1.2709 or CCZ is first solidified. In the context of 3D Multi-material fabrication, cleanliness means alien particles of Material A within the area of Material B or vice-versa.1550dACBFigure 4: Left: Representation of overlap C between two different materials A and B; Right: Representation of building order in case of 3D Multi-material partsResults and DiscussionMono-materialParameter analyses were performed and based on the need of high relative densities to reach good mechanical properties, the parameter set with the highest densities were used. Re-melting of the CCZ is necessary as a consequence of the limited laser power by 400 W and the goal of a good surface [12].

Re-melting of the CCZ is necessary as a consequence of the limited laser power by 400 W and the goal of a good surface [12]. The parameter are shown in Figure 5.Table 4 Optimized parameters with density results from the Archimedes method [12] MaterialHatch spacing (mm) 0.105 0.125Scan speed (mm/s) 600 600Laser power (W) Re-meltingRelative density (%) 100.00 ± 0.22 99.28 ± 0.251.2709 CCZ200 400No YesFigure 5 shows images of 1.2709 and CCZ made by a laser microscope. Both samples are producedusing the mentioned parameters.Figure 5: Images made by a laser microscope. Right: 1.2709 part etched with Nital. The arrow represents a melting line; Left: CCZ part etched with FeIIIClWhile the image of the 1.2709 shows clear visible the melting lines, they are not visible in CCZ.

The arrow represents a melting line; Left: CCZ part etched with FeIIIClWhile the image of the 1.2709 shows clear visible the melting lines, they are not visible in CCZ. This is due to the high heat conductivity of CCZ [15] and the re-melting strategy, which might cause a heat treatment- like process and therefore homogenizes the grain structure of the part.2D Multi-materialDuring the manufacture of 2D Multi-material parts, the focus was on the transition areas. Figure 6 shows an overview of one part per building order as well as a magnification of each transition area.1551Building-order CCZ – 1.2709 - CCZBuilding-order 1.2709 – CCZ – 1.2709Overview of entire sampleCCZCCZTransition 1.2709 - CCZ1.2709100µm1.2709100µm1.27091.2709Transition CCZ – 1.2709CCZ100µmCCZ100µmFigure 6: Images of sections of 2D Multi-material transitions in build-up direction. Pictures on the right for the transition area are etched.

Pictures on the right for the transition area are etched. Arrows point at crack within the steel area. The images of the transition area of the building-order 1.2709 – CCZ – 1.2709 are etched.The transition area for each kind of transition is similar in both building orders. In the case of CCZ- 1.2709, it is stricter and there is almost no mixture of the materials. This is in contrast to what occurs for the transition of 1.2709-CCZ. This has already been found in the case of laser welding of copper and stainless steel by Chen et al. [22]. They called it a “fusion welding mode” if there is a mixture of melt as a consequence of both materials being briefly in a liquid phase and therefore able to mix up. As a consequence of both materials being in a liquid state, fewer cracks occur within this kind of building order, since the liquid copper is also able to fill up the cracks. This is equivalent to melting copper on top of steel.

This is equivalent to melting copper on top of steel. A “welding-brazing-mode” occurs only if the steel is liquid, representing the welding part of the joint, and the copper remains in a solid phase, representing the brazing part of the joint. For 2D Multi-material LBM this is equivalent to fusing steel on top of copper.Focus on the etched images of the transition CCZ – 1.2709 shows melting lines in the copper area, although only in a small area above the transition. After this small area there are no more melting lines visible and the sample looks to the same as the images of the Mono-material process. In the 2D Multi-material process the differences in the thermal conductivity caused by the influence of 1.2709 might give the melting line more time for cooling and therefore make the melting lines visible.Mono-material with mixed powderBy building parts with mixed powder the influence of the material on the parameter and the resulting material composition can be determined.

The parameters were varied between the known set for the original powders represented in Table 4 corresponding to the material mixture. The parameters with the best relative densities are shown in Table 5. 1552Table 5 Parameters and relative density of samples with mixed powder Material 1.2709 (%) 25 50 75Hatch spacing (mm)Scan speed (mm/s)Laser power (W)CCZ (%) 75 50 250,115 0,118 0,125600 600 600300 333 400Re-melting Relative density(%)No No Yes99,96 99,98 99,99It can be seen that the parameters vary from the single-origin parameters because of the different material properties.

However, there is a bigger influence of the CCZ than the 1.2709, since the parameters, and therefore the fed energy, tends to be more similar to the parameters of CCZ.3D Multi-materialThe following Figure 7 shows 3D Multi-material samples built with different building-order andsolidification-order.Solidification-orderCCZ – 1.27091.2709 – CCZBuilding-orderCCZ1.2709Shell – Core1 mm1 mmCCZ1.2709Core – Shell1 mm1 mmFigure 7: Images of polished samples with varying solidification- and building-orderThere are visible differences in the quality of the parts depending on the solidification-order but no influences of the building-order. Since many cracks occur in the shell while first solidifying CCZ, this does not happen when 1.2709 is the first material solidified. There can be two reasons for this. First, corresponding to the results from the 2D Multi-material parts and from Chen et al.

There can be two reasons for this. First, corresponding to the results from the 2D Multi-material parts and from Chen et al. [22], both materials are in a solid phase while merging and therefore internal stresses in the material are reduced, if CCZ is solidifying second. The other possibility for the absence of cracks is that alien particles act as a source for cracks. That is why EDX-analyses were performed to analyze the distribution of Fe and Cu within the samples. These two elements were chosen since they represent the main part of each material. Results are shown in Figure 8.1553Solidification-orderCCZ – 1.27091.2709 – CCZBuilding-orderCCZ1.2709Shell – Core1 mm1 mmCCZ1.2709Core – Shell1 mm1 mmFigure 8: EDX analyses of 3D Multi-material samples with different building- and solidification order.

Fe is shown in blue and Cu in red: The black squares show the area of the coreIt can be seen that the 1.2709-powder can be removed more easily than the CCZ, since all images contain traces of Cu, while Fe is only in the area where it should remain – especially if this powder is delivered first. This supports the assumption that Fe-particles within the CCZ-part support the formation of cracks, since they might not become molten properly and therefore remain solid.It is conspicuous that in all cases the second delivered material is distributed over the entire sample. This can occur since the roughness of the first solidified material is too high and the second delivered powder fills up the gaps. The other possibility is that the first solidified material shrinks while solidifying [23] and the difference to the expected layer thickness gets filled up. In this case the cross-contamination will increase as the height of the part increases.

In this case the cross-contamination will increase as the height of the part increases. However, both assumptions will be analyzed in future investigations.Conclusion and OutlookA methodical approach was presented showing the procedure for building a 3D Multi -material part. After qualifying the CCZ and 1.2709 for the Mono-material process and therefore are able to build parts with a higher relative density of 99.5%. In a next step 2D Mutli-material parts were built showing the influence of the solidification-order on forming cracks in the steel. In the case of building CCZ on top of 1.2709 the quantity of cracks is decreased, since both materials are in liquid phase when CCZ becomes molten, since it requires more energy for melting, which leads to a mixture of both materials. The building of parts with mixed powder showed the influence of alien particles within the base material, since each mixture requires a different set of parameters for building parts with a high rel.

The building of parts with mixed powder showed the influence of alien particles within the base material, since each mixture requires a different set of parameters for building parts with a high rel. density. In future investigations this set of parameters will be used for the transition area to build parts with a specific material distribution in this area. In a final step, 3D Mutli- material parts were built using the parameters from the Mono-material process. Depending on the solidification- order, the cracks occurred within the steel area for the same reason as they did within the 2D Multi-material parts. Alien particles, especially of CCZ, were found within the other area by using EDX analyses.

Alien particles, especially of CCZ, were found within the other area by using EDX analyses. Two possible reasons – the shrinkage of the melting pool and the roughness of the first delivered material – will be investigated in future research.1554AcknowledgementsThe authors want to thank the Bavarian Research Foundation for the financial support within the project ForNextGen.References [1] J.-W. Choi, H.-C. Kim, R. Wicker, Multi-material stereolithography, Journal of Materials ProcessingTechnology 211 (2011) 318–328.[2] P. Muller, P. Mognol, J.-Y. Hascoet, Modeling and control of a direct laser powder deposition process forFunctionally Graded Materials (FGM) parts manufacturing, Journal of Materials Processing Technology 213 (2013) 685–692.[3] Y. Chivel, New Approach to Multi-material Processing in Selective Laser Melting, Physics Procedia 83(2016) 891–898.[4] C. Andriani, K.C. Chua, H.Z.

BRICATION OF 3D POLYMER-METAL NANO-COMPOSITES IN A SINGLE STEP BY TWO-PHOTON INDUCED POLYMERISATION AND METAL SALT REDUCTIONQin Hu*, Yaan Liu, Yinfeng He, Fan Zhang, Ricky Wildman, Chris Tuck and Richard HagueEPSRC Centre for Innovative Manufacturing in Additive Manufacturing Faculty of Engineering, University of Nottingham Nottingham, UK, NG7 2RDEmail: qin.hu@nottingham.ac.ukAbstractFabrication of 3D polymer-metal nano-composites in a single step by two-photon induced polymerisation and metal salt reduction has been demonstrated in this study. Two kinds of composites, SU8-Au and IPL-Au, based on different mechanisms of polymerisation, have been fabricated and compared. To the best of our knowledge, this is the first demonstration of IPL-Au nano-composites being fabricated by two-photon lithography. Extra photoninitator is needed for the fabrication of IPL-Au composites, to provide extra free radicals to sustain the two reactions processing at the same time.

Extra photoninitator is needed for the fabrication of IPL-Au composites, to provide extra free radicals to sustain the two reactions processing at the same time. The distribution of the generated Au nanoparticles in IPL matrix is more uniform than that in SU8 matrix. The technique demonstrated in this study can have great application in metamaterial fabrication.KeywordsNano-composites, polymerisation, metal salt reduction, metamaterialstwo-photonlithography,goldnanostructures,nanofabrication,1. IntroductionMetamaterials are novel artificial multi-material constructs with sub-wavelength-sized periodic structures containing both conducting (metal) and dielectric (polymer) components. Their unique but previously unachievable properties have promoted lots of extraordinary applications (Cai, Chettiar et al. 2007; Engheta 2007; Choi, Lee et al. 2011; Soukoulis and Wegener 2011).

Their unique but previously unachievable properties have promoted lots of extraordinary applications (Cai, Chettiar et al. 2007; Engheta 2007; Choi, Lee et al. 2011; Soukoulis and Wegener 2011). The current techniques for metamaterials manufacturing are mainly based on traditional microchip patterning techniques, e.g. photolithography combined with metal deposition, that are expensive, multistep and slow (Liu, Zhu et al. 2012; Moser and Rockstuhl 2012). Two-photon lithography, which uses a focused ultrafast laser beam to penetrate into photo-reactive polymer resin and scans in true 3D, can overcome this limitation and fabricate micro/nano structures in arbitrary 3D with around 100 nm resolution in a fast and cost- efficient way (Lee, Yang et al. 2006; Maruo and Fourkas 2008; Park, Yang et al. 2009; Malinauskas, Farsari et al. 2013).

2006; Maruo and Fourkas 2008; Park, Yang et al. 2009; Malinauskas, Farsari et al. 2013). Both polymeric structures and metallic structures have been manufactured respectively by two-photon induced polymerisation and metal salt reduction (Tanaka, Ishikawa et al. 2006; Kaehr and Shear 2008; Seidlits, Schmidt et al. 2009; Xu, Xia et al. 2010; Greiner, Richter et al. 2012; Ishikawa and Tanaka 2012). Early work of1036fabricating metallic structures in a polymeric matrix by two-photon associated processing has been demonstrated by Prasad et al (Shukla, Furlani et al. 2010; Shukla, Vidal et al. 2011; Vora, Kang et al. 2012). The great potential applications urge us to expand current capabilities and especially in enriching materials available for various needs.The mechanism of photopolymerisation can be generally divided into two groups – free radical-based polymerisation and ionic-based polymerisation (Lee, Kim et al. 2008).

2008). In the current study, two-photon induced polymerisation based on both mechanisms has been compared, using two commercial resins – SU8 and IPL. SU-8 is a commonly used epoxy- based resin, and the polymerisation is based on the ionic mechanism. The commercially available SU8 resin contains a small amount of photoinitiator(s), optimized for traditional UV lithographic applications. The contained initiator(s) may vary due to different suppliers. IPL is a free radical-based commercial photo-reactive resin, developed by Nanoscribe GmbH for two-photon polymerisation. The as-purchased IPL resin contains <5% initiator. We demonstrate in this paper that two kinds of polymer-metal nano-composites – SU8-Au and IPL-Au – can be fabricated in a single step by two-photon induced polymerisation and metal salt reduction. To the best of our knowledge, this is the first demonstration of IPL-Au nano- composites being fabricated by two-photon lithography.2.

To the best of our knowledge, this is the first demonstration of IPL-Au nano- composites being fabricated by two-photon lithography.2. ExperimentalNano-composites were fabricated using a commercial two-photon lithography system – Nanoscribe Photonic Professional GT. The system uses a fiber laser at a wavelength of 780 nm, pulse frequency of 80 MHz and pulse duration of 120 fs. Laser beam is focused by an oil immersion objective (63x, NA = 1.4, WD = 190 µm). Micro/nano patterns are formed by moving the laser beam using a high speed X-Y galvo-scanner or shifting the sample position using a piezo X-Y-Z stage. All the patterns shown in this paper are formed by using the galvo mode at a writing speed of 10,000 µm/s and the laser power in the range of 25 mW to 50 mW.Gold chloride hydrate powder (HAuCl4∙xH2O) purchased from Sigma-Aldrich was used as source for metallic component. Monomer SU8 and IPL were purchased from Gersteltec and Nanoscribe, respectively.

Monomer SU8 and IPL were purchased from Gersteltec and Nanoscribe, respectively. Photoinitiators were sourced from Sigma-Aldrich.Gold chloride powder was first dissolved in solvent before mixing with monomer resin. Ethanol was used in the mixture of IPL with Au, and 1,4-Dioxane was used in the mixture of SU8 with Au. Then a drop of mixture was loaded on a clean cover slip for two- photon processing. After laser exposure, the cover slip was soaked in a suitable solvent (developer) to remove the unsolidified resins. The fabrication procedure is illustrated in Figure 1.Structures were imaged by a desktop Scanning Electron Microscopy (SEM) system - Hitachi TM3030. The chemical compositions were analysed by an Energy-dispersive X-ray Spectroscopy (EDX) system - Bruker Quantax 70 - integrated within the Hitachi SEM.1037Figure 1: Illustration of fabrication steps.3.

Results and discussionFabricating nano-composites in a single step by two-photon induced polymerisation and metal salt reduction is illustrated in Figure 2. A near-infrared femtosecond laser beam is focused into a photo-reactive resin containing the mixture of monomer and metal salt. A photoinitiator is excited by the simultaneous absorption of two photons, which can trigger local chemical reactions, including monomer cross linking and metal salt reduction. In this way, both polymer and metal nanoparticles can be formed. Due to high surface tension, the generated metal nanoparticles tend to aggregate into large ones. The local monomer cross linking may serve as a matrix, thus reduce the diffusion-related metal aggregation.

The local monomer cross linking may serve as a matrix, thus reduce the diffusion-related metal aggregation. These two simultaneous chemical reactions may compete with each other, leading to different morphology.Figure 2: Illustration of fabricating nano-composites in a single step by two-photon induced polymerisation and metal salt reduction1038Nano-composite of SU8 and AuThe mixture of gold chloride and SU8 was exposed to laser at energy of 30 mW and a scanning speed of 10,000 µm/s. A SU8-Au hybrid pattern is shown in Figure 3. The grey lines were made of SU8, and the white dots were Au nanoparticle aggregations, with the sizes ranging from dozens of nanometers up to 1 micrometer. The mechanism of simultaneous Au salt reduction and SU8 ionic polymerization has been discussed before (Shukla, Furlani et al. 2010; Shukla, Vidal et al. 2011). Compared to the patterns produced by Prasad et al (Shukla, Furlani et al. 2010; Shukla, Vidal et al.

2010; Shukla, Vidal et al. 2011). Compared to the patterns produced by Prasad et al (Shukla, Furlani et al. 2010; Shukla, Vidal et al. 2011), the aggregation of Au nanoparticles at some positions is more predominant for our case. This phenomenon could be associated with the different photoinitiators used. Our SU8 sourced from Gersteltec contains no antimony and we have not added any extra initiators for the current study, while the SU8 used by Prasad et al -). As the exact initiators which was sourced from Microchem includes antimony (Ar2I+SbF6 in our SU8 are not clear, it is difficult to identify how they affected the aggregation of in-situ generated nanoparticles.Figure 3: SEM image of SU8-Au nano-composite. The grey lines were made of SU8, and the white dots embedded in the grey lines were Au nanoparticle aggregations.Nano-composite of IPL and AuThe following study was firstly carried out by testing the sample containing IPL and gold chloride without adding extra initiators.

Although commercial IPL is designed for two-photon processing, no polymerisation was observed for the sample containing the mixture. As shown in Figure 4, only Au nanoparticle aggregation is observed. This is because two-photon induced metal ion reduction consumes free radicals generated by initiators. However free radicals are also needed for two-photon induced polymerisation. In the case of lacking sufficient free radicals, part of reactions is suppressed.

However free radicals are also needed for two-photon induced polymerisation. In the case of lacking sufficient free radicals, part of reactions is suppressed. In current situation, gold ions reduction consumed all the free radicals in the mixture, thus prevented monomers from cross- linking.1039(a) (b) Figure 4: (a) SEM image showing the fabricated pattern using the mixture of IPL with gold chloride without extra initiators; (b) Magnified SEM image showing the section marked in figure (a).Extra photoinitiator was added to the mixture of IPL and gold chloride, to provide extra free radicals, thus to promote the two reactions – polymerisation and reduction – to happen at the same time. The initiator was chosen due to its high two-photon absorption cross-section at the wavelength of 780 nm - the wavelength of the laser used in this study (Makarov, Drobizhev et al. 2008).

The initiator was chosen due to its high two-photon absorption cross-section at the wavelength of 780 nm - the wavelength of the laser used in this study (Makarov, Drobizhev et al. 2008). Figure 5(a) shows a 10-µm-high micro structure formed by adding a small amount of initiator to the mixture of IPL and gold chloride. The surface morphology of the fabricated micro structure looks no different from the structure formed by using pure IPL, i.e., containing no metallic part. However, EDX analysis shows that Au nanoparticles are well embedded in the IPL matrix. As shown in Figure 5(b), the distribution of Au element corresponds well with the 3D micro structure. However, the distribution of Cl is almost uniform throughout the whole substrate (see Figure 5(c)). EDX analysis proves that gold chloride has been successfully reduced into gold.Au Cl(a) (b) (c) Figure 5: (a) SEM image of IPL-Au nano-composite fabricated using rhodamine B as an extra photoinitator.

The pattern is about 10 µm high, with 25 layers. (b) & (C) EDX mapping analysis showing Au and Cl distribution in the area marked in figure (a), which proved that gold chloride has been successfully reduced into gold.Comparing the two samples shown in Figure 3 and 5, it is obvious that Au nanoparticles distributed more uniformly in the matrix of IPL than that in the matrix of SU8.1040The aggregation of Au nanoparticles in SU8 matrix is also observed in the work published by Prasad et al (Shukla, Vidal et al. 2011). Such great difference in surface morphology can be associated with the different polymerisation mechanism. Ionic-based polymerisation (for the case of SU8) tends to generate more heat than radical-based polymerisation (for the case of IPL). The local heat can accelerate the surface diffusion of Au nanoparticles, thus promote aggregation. To fabricate nano-composite with high resolution, radical-based polymerisation is preferred.4.

The local heat can accelerate the surface diffusion of Au nanoparticles, thus promote aggregation. To fabricate nano-composite with high resolution, radical-based polymerisation is preferred.4. ConclusionsTwo kinds of polymer-metal nano-composites – SU8-Au and IPL-Au – have been fabricated in a single step by two-photon induced polymerisation and metal salt reduction. Extra photoinitiator is needed for the generating the composite of IPL-Au, to provide extra free radicals to sustain the two reactions processing at the same time. The distribution of generated Au nanoparticles in IPL matrix is more uniform than that in SU8 matrix.

brication of Advanced Thermionic Emitters Using Laser Chemical Vapor Deposition-Rapid PrototypingBrian T. Fuhrman, Chad E. Duty, Daniel L. Jean, W. Jack Lackey Rapid Prototyping and Manufacturing Institute Woodruff School of Mechanical Engineering Georgia Institute of Technology Atlanta, GA 30332-0405AbstractLaser Chemical Vapor Deposition-Rapid Prototyping (LCVD-RP) is a relatively new manufacturing process. Its capabilities are ideally suited for the manufacturing of a type of electron emitter called an integrated-grid thermionic emitter. The integrated-grid thermionic emitter is composed of wagon wheel-like structures of alternating layers of boron nitride and molybdenum on tungsten. The goal of this paper is to determine the feasibility of using LCVD- RP technology to manufacture advanced thermionic emitters.IntroductionThe purpose of this paper is to determine the feasibility of using LCVD-RP to produce a thermionic emitter used in a Pierce-type electron gun.

The Pierce-type electron gun utilizes a thermionic emitter composed of wagon wheel-like grid structures of molybdenum suspended over a concave piece of barium impregnated tungsten. The performance of the electron gun is dependent on minimizing the distance between the tungsten and the molybdenum grid. Current manufacturing methods mechanically suspend the grid over the tungsten. Although this method is currently more cost-effective, it is the limiting factor for the performance of the emitter. LCVD-RP has the potential to manufacture advanced thermionic emitters by reducing the distance between the metal layers and the diameter of the emitter. However, the potentially high cost of developing and implementing this technology may not economically justify the improvements in performance.LCVD is the process of using the thermal energy from a laser to induce a chemical reaction to occur within special reagent gases to produce a solid deposit.

Depending on the reagent gases, a variety of metals and ceramics can be precisely deposited onto a substrate. Since the material is deposited on the atomic level, the deposit is fully dense and homogenous. The material properties of the deposit more closely match the theoretical bulk properties. Current metal and ceramic rapid prototyping processes, such as selective area sintering, sinter metal or ceramic powder particles together. The properties of the material are dependent on the strength of the bond between the sintered particles.Much work has been done in the field of LCVD in the past 25 years. A wide variety of metals and ceramics such as Ag, B, Si, C, SiC, Si3N4, Al2O3, TiN, TiC, and Ni have been successfully deposited as fibers utilizing LCVD.1,2 In addition to fibers, several researchers have been successful in depositing materials in the form of films and lines.

Mutoh et al.3 deposited498alternating layers (5 to 25 nm thick) of tungsten and carbon by quickly scanning a laser over a relatively large cross-section. Other researchers have been successful in the direct writing of TiN4, Al5, Au6, Si7, GaAs8, and Cu9 onto various substrates.In the past 10 years, several researchers were successful in depositing three-dimensional structures using LCVD-RP. In 1991, Zong10 created a rectangular block of graphite. Other shapes such as a sphere, an hourglass, and a micron-sized needle were created from C, Ni, W, and Fe.11,12 Maxwell et al. 13,14 deposited carbon in the form of helical springs. Fiber diameters of the spring as small as 5 m m and spring internal diameters as small as 60 m m were achieved. These springs were utilized for functioning microsolenoids.LCVD-RP Thermionic EmittersThermionic emitters are used as the electron source inside electron guns.

These springs were utilized for functioning microsolenoids.LCVD-RP Thermionic EmittersThermionic emitters are used as the electron source inside electron guns. Devices such as television tubes, microwave emitters, oscilloscopes, and cathode ray tubes utilize an electron gun to create the picture on the screen. A thermionic emitter is composed of a piece of metal called a cathode with a metallic control grid suspended over it. Electrons are produced by the thermionic emitter by applying a high negative voltage to the cathode (generally, the cathode voltage is referenced to zero). The number of electrons that are allowed to pass are controlled by the voltage applied to the control grid. The closer the voltage on the control grid is to the cathode voltage, the greater the number of electrons are repelled back to the cathode.The accuracy and precision of thermionic emitters is limited by current manufacturing techniques.

The control grid has to be mechanically supported above the cathode by means of rods or other mechanical devices. If the grid is too close to the cathode, the heat from the cathode can warp the control grid causing imprecise operation and even a short circuit. The closer the control grid is to the cathode, the better the sensitivity of the thermionic emitter.Fabricating an integrated grid design where the cathode and control grid are one unit is an attractive alternative. However, current manufacturing methods that would apply in this situation are limited to planar assemblies. Ohlinger15 suggested that LCVD-RP could be used to create miniature, stable, high power bonded grid emitters . He suggested that a wagon wheel-like grid structure be created on a concave tungsten substrate as illustrated in Figure 1. The grid would consist of alternating layers of boron nitride and molybdenum.

The grid would consist of alternating layers of boron nitride and molybdenum. Because of the properties of boron nitride, the boron nitride layers would isolate the molybdenum grid from the tungsten substrate and also prevent their distortion. By using the boron nitride layers, the grid can be physically closer to the cathode without the problems associated with previous designs.Figure 1: Thermionic emitter via LCVD-RP.15499Experimental DesignThe goal of this research was to determine the feasibility of using LCVD-RP to create a thermionic emitter as illustrated in Figure 1. If alternating lines of molybdenum and boron nitride could be deposited onto a tungsten substrate, this process could be used to create a thermionic emitter. Experimentation was done on depositing boron nitride onto both tungsten sheet (99.95% pure, 0.01 inch thick) and molybdenum sheet (99.95% pure, 0.05 inch thick).

Experimentation was done on depositing boron nitride onto both tungsten sheet (99.95% pure, 0.01 inch thick) and molybdenum sheet (99.95% pure, 0.05 inch thick). To acquire the process parameters for depositing molybdenum, 0.75-inch diameter, 0.125-inch thick graphite disks (ATJ grade graphite from UCAR Carbon Company) were used as substrates.Boron nitride can be deposited from boron trichloride, ammonia, and hydrogen via thefollowing equation:BCl3(gas) + NH3(gas) ﬁBN(solid) + 3HCl(gas)Hydrogen gas (H2) was used to dilute the mixture and to react with the chloride to eliminate all of the BCl3 (hydrogen was not included in the above chemical reaction). Because of the vapor pressure of BCl3, it was necessary to heat both the BCl3 line and the other reagent line to approximately 50°C. As has been reported,16-18 BCl3 and NH3 can react at lower temperatures to produce a white powdery substance of NH4Cl.

As has been reported,16-18 BCl3 and NH3 can react at lower temperatures to produce a white powdery substance of NH4Cl. To reduce the formation of this powder on the substrate, the substrate was globally heated to approximately 350°C. In addition, the BCl3 and NH3 reagents were allowed to mix approximately 15 to 30 mm from the substrate (more detail is given in the Experimental Setup section).The deposition of molybdenum from molybdenum pentachloride and hydrogen isillustrated by the following reaction:MoCl5(gas) + 5/2H2(gas) ﬁMo(solid) + 5HCl(gas)It was necessary to sublimate MoCl5 into a gas because it is a solid powder at room temperature and atmospheric pressure. Saeki et al.19 experimentally related the sublimation temperature to the partial pressure of MoCl5 by the following equation:log( PM)=3504 T+.9465Where Pm is the partial pressure of MoCl5 in torr and T is the sublimation temperature in Kelvin.

This equation is applicable between 80 and 150(cid:176) C (353 and 423 K).Experimental SetupThe design and capabilities of the experimental system used to conduct this research have been described in detail elsewhere.20,21 As an overview for this paper, a brief description of the system follows. The primary components of the experimental system consists of the reaction500chamber, the laser, the reagent and vacuum lines, and the laser beam delivery system. Figure 2 is an illustration of the functions of the reaction chamber.The equipment was actually composed of two separate chambers. The upper chamber (as illustrated in Figure 2) was the actual reaction chamber. The lower chamber was responsible for housing and protecting the three movable stages from the reagent gases (not shown). The upper several novel chamber contained features the deposition in process. To improve the deposition rate, the reagent gases were delivered to the reaction zone via a jet nozzle.

The upper several novel chamber contained features the deposition in process. To improve the deposition rate, the reagent gases were delivered to the reaction zone via a jet nozzle. To reduce the deposition temperature the laser had to achieve, a resistive heater was used to globally heat the substrate. Other features included the use of a CCD camera to visually monitor the process and a distance- the measuring device height of the deposit (this feature was not used for these experiments). A 100-Watt, CO2 laser focused approximately to 200 m m, irradiated the substrate.Focusing LensDistance DeviceLaser BeamCCD Camera Portto aidDual NozzleSubstrateDepositHeaterVacuum Portto measureFigure 2: Schematic of LCVD-RP reaction chamber.For the deposition of boron nitride, it was discovered that if the reagents were allowed to mix before the chamber, a white, powdery substance appeared within the chamber.

Therefore, concentric nozzles with BCl3 in the outer nozzle and the other reagents in the inner nozzle were used to lessen the mixing distance from the substrate. The nozzle was approximately 15 mm from the surface of the substrate.Thermodynamic ModelA thermodynamic model of the interaction between the reagents used for depositing boron nitride and the tungsten or molybdenum substrate was developed. Deposition temperatures from 300 to 1500 K, deposition pressures of 0.5 and 1.0 atmosphere, and NH3 to BCl3 ratios of 1 to 5 were used as input conditions for the model. Because of the high deposition temperatures for boron nitride, there exists the possibility that the molybdenum or tungsten could have reacted with the chlorine from BCl3 to produce a molybdenum or tungsten chloride.

After the initial experimentation, a white powdery substance that was believed to be NH4Cl (to be discussed in further detail in the Results and Conclusions section) was deposited on the surface of the substrate. Because the majority of the substrate and surrounding chamber was close to room temperature, the thermodynamic model of the substrate was carried out at the lower temperatures to confirm the experimental results.501The computer program SOLGASMIX-PV was used to analyze the thermodynamic reactions. SOLGASMIX-PV utilizes an iterative method to determine the minimum Gibb’s free energy of the mixture.

SOLGASMIX-PV utilizes an iterative method to determine the minimum Gibb’s free energy of the mixture. The change in Gibb’s free energy is related to the entropy and enthalpy of the elements and compounds by the following equation:DD-DSTHGiiiWhere D Gi is the change in Gibb’s free energy (kJ/mole) for the ith species, D Hi is the change in enthalpy (kJ/mole) of the ith species, T is the temperature in Kelvin, and D Si is the change in entropy (kJ/mol/K) of the ith species. The calculations were dependent on the quantities of the reactants, the temperature and pressure of the system, and the respective enthalpy and entropy of the possible products (which can be solid, liquid, or gas).

Confirming the experimental results, the computer program calculated that NH4Cl formed in addition to solid BN at reaction temperatures from 300 to 600 K. Figure 3 is a graph of the moles of NH4Cl formed versus the deposition temperature. Each line represents the amount of NH3 used in The the thermodynamic calculations. calculations indicated that if the NH3 to BCl3 ratio was one to one, only a small Each fraction of NH4Cl is formed. additional mole of NH3 used resulted in the formation of an increasingly greater quantity of NH4Cl.502ResultsThe deposition of boron nitride lines from BCl3 and NH3 was not successful. Instead, the only material to be deposited was some form of boron oxide and what is believed to be NH4Cl. During the experiments, a white film that was believed to be NH4Cl formed on the surface of the substrate. This hypothesis is plausible due to the lower temperature of the substrate and the surrounding reaction chamber.

This hypothesis is plausible due to the lower temperature of the substrate and the surrounding reaction chamber. Experiments were conducted with the substrate globally heated to approximately 620 K with the same results.Only a small amount of molybdenum was successfully deposited using MoCl5. Figure 4 is an SEM micrograph and EDS analysis of one of the attempts at depositing a line of molybdenum. This deposit was created at a deposition pressure of 5.0 psi, laser power of approximately 75 Watts, MoCl5 flow rate of 6.22 sccm, sublimation chamber temperature of 150(cid:176) C, and scan speed of 0.02 in/min. The partial pressure of the MoCl5 was calculated to be 15 torr. The EDS analysis indicated that only a trace amount of molybdenum was evident on the substrate. The presence of oxygen was due to a bad seal on the sublimation chamber.

The EDS analysis indicated that only a trace amount of molybdenum was evident on the substrate. The presence of oxygen was due to a bad seal on the sublimation chamber. The reagent powder had been contaminated with outside air.a)600 m mFigure 4: a) SEM micrograph and b) EDS analysis of the line produced by the attemptat depositing molybdenum.Conclusions and RecommendationsNeither boron nitride nor molybdenum were successfully deposited in these experiments. However, LCVD-RP is still a possible manufacturing tool for depositing the structure necessary for the integrated-grid thermionic emitter. Because of the production of the by-product, NH4Cl, depositing boron nitride from BCl3 and NH3 is not suited for this experimental system. Even though the substrate can be globally heated past the threshold temperature for the formation of NH4Cl, the remainder of the reaction chamber is still at room temperature.

Even though the substrate can be globally heated past the threshold temperature for the formation of NH4Cl, the remainder of the reaction chamber is still at room temperature. The use of other reagents that do not require NH3, such as B3N3H3Cl3 (known as trichloroborazine or trichloroborazol),23,24 would eliminate the formation of NH4Cl or any solid by-product.503b)Trace amounts of molybdenum were deposited using this experimental system. Due to the contamination of the reagent with air, the feasibility of using MoCl5 as a reagent for depositing molybdenum is uncertain. In addition, no experimentation was attempted on depositing molybdenum onto previously deposited boron nitride. The morphology and characteristics of the deposit could be different on boron nitride than on graphite.The manufacture of integrated-grid thermionic emitters via LCVD-RP is still a viable option. By using a different reagent for boron nitride, the formation of solid by-products could be eliminated.

By using a different reagent for boron nitride, the formation of solid by-products could be eliminated. Further experimentation would be required to prove or disprove the use of MoCl5 to deposit molybdenum.AcknowledgementsThe authors would like to thank Lisa Detter-Hoskin and Dr. Paul Kohl for the use of theSEM equipment and the glove box, respectively.References1.2.3.4.5.6.7.8.9.L. S. Nelson and N. L. Richardson, “Formation of Thin Rods of Pyrolytic Carbon by Heating with a Focused Carbon Dioxide Laser,” Material Research Bulletin, 7, 971- 976, 1972. D. Bauerle, Chemical Processing with Lasers, Springer-Verlag: Berlin, 1986, 224, 71- 93. K. Mutoh, Y. Yamada, T. Iwabuchi, and T. Miyata, “Thickness Control of Multilayer Films in Laser-Induced Chemical Vapor Deposition,” Japanese Journal of Applied Physics, 31, 1183-1184, 1992. A. Silvestre, M Parames, and O.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFabrication of aligned nanofibers along Z-axis – A novel 3D Electrospinning techniqueGeorge Z Tan1*, Yingge Zhou11Department of Industrial, Manufacturing, and Systems Engineering Texas Tech University, Lubbock, TX *corresponding author: george.z.tan@ttu.eduAbstractThis study presents a 3D fabrication technique of nanofibrous scaffold for tissue engineering. A divergence static electric field was introduced in an electrospinning system to induce a self- assembly of aligned nanofibers into a tunable 3D architecture with thickness ranging from 2-12 mm. The effects of collector configuration on polycaprolactone (PCL) nanofiber attributes were investigated. Human fibroblast cells were cultured on the nanofiber scaffold in vitro for 7 days.

The effects of collector configuration on polycaprolactone (PCL) nanofiber attributes were investigated. Human fibroblast cells were cultured on the nanofiber scaffold in vitro for 7 days. It was found that the width and inclination angle of the collector influenced the nanofiber density distribution. The cells proliferated on the scaffold and organized as a fibrous matrix which mimicked the microstructure of native musculoskeletal tissues.Keywords: divergence electrospinning, 3D nanofiber scaffold, tissue engineeringIntroductionThe demand for human organ substitutes and the limited availability of transplant donors has advanced the development of tissue engineering. Tissue engineering generally requires the use of artificial scaffolds to recreating the in vivo milieu which coordinates a combination of biophysical and biochemical stimuli governed by complex spatiotemporal mechanisms to control cell behavior [1, 2].

The scaffold fabrication processes can generally be classified as top-down or bottom-up methods [3]. In top-down approaches, the biomaterials may be shaped to resemble biological geometries first, and the cells then seeded homogenously in the scaffold. With bottom-up approaches, modular units of cells and biomaterials are integrated to form 2D or 3D composite architecture. Common top-down approaches include solvent casting, particulate-leaching techniques, gas foaming, phase separation, freeze drying, melt molding, electrospinning and others. Research in this area tends to focus on optimization of the scaffold design by controlling the pore characteristics and reinforcing the biological properties. Bioprinting is one of the most extensively studied bottom-up techniques for scaffold development. In bioprinting, small units of cells and biomaterials are dispensed with high precision to form tissue-like structures in a layer-by-layer fashion.

In bioprinting, small units of cells and biomaterials are dispensed with high precision to form tissue-like structures in a layer-by-layer fashion. The three major bioprinting techniques are inkjet, laser-assisted, and extrusion [4]. To date, no single technique enables the production of all scales and complexities of synthetic tissues [3].Recent bioinspired approaches focus on creating a biomimetic cell microenvironment closely resembling the natural gradients in cell distribution, extracellular matrix (ECM), and tissue topology. ECM is a collection of extracellular molecules secreted by cells that provides structural and biochemical support to the surrounding cells [1]. Gels of polysaccharides and fibrous proteins fill the interstitial space and act as a compression buffer against the stress placed on the ECM [2].

Gels of polysaccharides and fibrous proteins fill the interstitial space and act as a compression buffer against the stress placed on the ECM [2]. Studies show that micro- and nano-topography and local environment of ECM influence trends in cell behavior by providing biochemical and biophysical stimuli to promote cell adhesion,2396proliferation, morphogenesis, and motility [3, 4]. Critical physical features of ECM include its dimensionality, architecture, stiffness, ligand topography, and density [5]. Despite rapid progress in engineering relatively simple tissues, there is a research gap in integrating the tunable microarchitecture in heterogeneous scaffold development to closely resemble natural tissue structures.

Especially, the dimension constraints limit the further applications of existing additive manufacturing techniques to mimic the unique structures and properties of natural tissues.In this paper we presented a rapid fabrication technique of 3D nanofibrous scaffold for tissue engineering. A divergence static electric field was introduced in an electrospinning system to induce a self-assembly of aligned polycaprolactone (PCL) nanofibers into a 3D architecture. A cell culture test was performed to demonstrate the feasibility of providing 3D physical cues to guide cell growth by the nanofiber scaffold.Materials and Methods2.1. Polymer solution preparationPCL polymer solution was prepared by dissolving 15 wt.% PCL pellets (MW = 80,000, Sigma- Aldrich®, St. Louis, MO) in organic solvent. The solvent was prepared by homogenizing N,N- Dimethylformamide (99.8%, Sigma-Aldrich®, St. Louis, MO) and acetone (Thermo Fisher Scientific, Waltham, MA) at volume ratio of 1:1.

The solvent was prepared by homogenizing N,N- Dimethylformamide (99.8%, Sigma-Aldrich®, St. Louis, MO) and acetone (Thermo Fisher Scientific, Waltham, MA) at volume ratio of 1:1. The solution was magnetically stirred for 4 hours at room temperature until there was no visible suspended particulates. To stain the nanofibers, 0.1 mg/ml of DilC12(3) perchlorate (Thermo Fisher Scientific, Waltham, MA) was added to the PCL solution. All materials were purchased from Sigma-Aldrich® (St. Louis, Missouri, United States).2.2. Electrospinning configurationThe divergence electrospinning was configured based on a high voltage direct current power source and a specially-designed polylactic acid (PLA) collector. The design of the collector was demonstrated in Figure 1a. The collector was comprised of two tilted surfaces, which had an inclination angle of 115 degrees to the collector bottom. This double-bevel collector was 3D printed by the Ultimaker3 system (Ultimaker®, Netherlands).

This double-bevel collector was 3D printed by the Ultimaker3 system (Ultimaker®, Netherlands). Aluminum foil was glued on the inner surface of each bevel using PCL solution, and was connected to the aluminon wire through a small central hole on the bevel.2397Figure 1. (a) Collector design. (b) Divergence electrospinning configurationThe collector was placed on an insulating stand in the electrospinning chamber. The aluminum foils on collectors were grounded separately using electrical wires passing through the inner cavity of the stand (Figure 1b). The PCL solution syringe was set to be 10 cm above the bottom center of the collector, and was pumped with a rate of 1 mL/hr. The voltage was set to be 10 kV, and the electrospinning time was approximately 2 minutes. The double-bevel induced a divergent electric field so that nanofibers were deposited onto both bevels of the collector while formed connecting fibers in between.2.3.

The double-bevel induced a divergent electric field so that nanofibers were deposited onto both bevels of the collector while formed connecting fibers in between.2.3. Scaffold characterizationTo characterize the internal nanofiber distribution, the scaffold was vertically sectioned by glass slides as illustrated in Figure 2. The top and bottom sectioning was performed at the 2 mm from the top and bottom surface, respectively. The nanofiber morphology at each section was examined by scanning electron microscopy (SEM, Hitachi S-4300). Fiber diameter, density and alignment were analyzed by ImageJ after the SEM images was segmented and processed. Fiber density was characterized by counting the number of fibers that intersected a line drawn across the middle of a binary SEM image. Fiber alignment were characterized by convergence degree of the fiber orientation histogram.2398Figure 2. Sectioning of the nanofiber scaffold.2.4.

Fiber alignment were characterized by convergence degree of the fiber orientation histogram.2398Figure 2. Sectioning of the nanofiber scaffold.2.4. Cell CultureWe fabricated a scaled-down scaffold (20 mm × 10 mm, width × length) to fit the tissue culture plate. The collector was immersed in 70% ethanol for 30 minutes prior to electrospinning. The electrospinning chamber was sprayed and wiped with 70% ethanol and maintained sterile during electrospinning process. After the nanofiber scaffold was obtained, the collector and the scaffold were sterilized by UV light in the biosafety cabinet for 30 minutes. The scaffold in the collector was then placed in a sterile 50 ml beaker. Human fibroblast cells (ATCC®, MRC-5) were suspended in complete growth medium. The medium was slowly added into the beaker to cover the scaffold. The cell-seeded microfiber was incubated for 7 days. Cells were fixed with 4% formaldehyde after 7 days.

The medium was slowly added into the beaker to cover the scaffold. The cell-seeded microfiber was incubated for 7 days. Cells were fixed with 4% formaldehyde after 7 days. For microscope imaging, cells were stained with CruzFluor Phalloidin fluorescence conjugate and 4',6-diamidino-2-phenylindole (DAPI) for cell cytoplasm and nucleus, respectively.ResultsThe divergent electric field induced a nanofiber deposition on the surfaces of both bevels, meanwhile the whipping process led to a self-assembly of nanofiber between the two bevels. As a results, an aligned fiber bundle structure gradually formed between the two bevels from the bottom to the top (Figure 3). The thickness of the obtained nanofiber scaffold was approximately 20 mm. Fluorescent nanofibers at the top and bottom sections were presented in Figure 4. It was observed that the fiber density was higher at the top layers than at the bottom layers.Figure 3. Formation of the 3D nanofiber scaffold.

It was observed that the fiber density was higher at the top layers than at the bottom layers.Figure 3. Formation of the 3D nanofiber scaffold. (a)-(c) represent 10 secs, 30 secs, and 60 secs after electrospinning, respectively.Figure 4. Fluorescent nanofibers at (a) the top section, and (b) the bottom section of the scaffold. Scale bar = 200 µm.2399A summary of fiber diameter distribution and scaffold porosity was shown in Figure 5. The diameter of the nanofibers ranged from 200 nm to 780 nm. The mean values were was 468 nm at the top section, and 514 nm at the bottom section. No substantial difference in fiber diameter was found. On the other hand, a gradient in scaffold porosity along the z-axis of the collector was observed. The collector saw the densest fibers on both top and peripheral areas, resulting in a low porosity, while less fibers were formed at the bottom and central areas, resulting in a high porosity.Figure 5.

A summary of nanofiber diameter and scaffold porosity.Representative images of cell culture were shown in Figure 6. The cells stretched along the nanofibers after 24 hours, and proliferated in the scaffold during the 7 days. The florescent image (Figure 7) showed the cell skeleton and organization, proving that the nanofiber scaffold guided the cell alignment and growth.Figure 6. Fibroblast cells seeded on the scaffold after (a) 2 hours and (b) 24 hours.2400Figure 7. Florescent cells after 7 days. (a) nucleus stained by DAPI, (b) cytoplasm by Phalloidin, and (c) anoverlap.DiscussionAs the ECM is a composite structure combining fibrous proteins held within gelatinous ground substance, the integration of micro-/nano-fibers and hydrogels may offer potential for development of superior composite scaffolds resembling the natural 3D environments with enhanced mechanical properties and cell penetration [6].

This becomes especially critical in applications involving musculoskeletal soft tissues such as tendons, ligaments, knee menisci, and the annulus fibrosus (AF) of the intervertebral disc, which have a more pronounced fibrous cytoskeletal organization [7]. Although the concept of combining hydrogel and microfiber is in its infancy, a number of pioneering works have been explored to create biomimic scaffold for different applications.One of the most straightforward methods of creating fiber-embedded hydrogel is to mix cell- embedded medium with microfibers and then crosslink the hydrogel. Narayana et al. 3D-printed alginate hydrogel containing polylactic acid (PLA) nanofibers and human adipose-derived stem cells [7]. The electrospun PLA fibers were bended in alginate by vortex and sonication. The composite hydrogel was 3D printed into a shape of human medial knee meniscus.

The electrospun PLA fibers were bended in alginate by vortex and sonication. The composite hydrogel was 3D printed into a shape of human medial knee meniscus. Results showed that the nanofiber-reinforced bioink allowed higher levels of cell proliferation within bioprinted strands. Fiber-hydrogel composites can also be achieved by encapsulating pre-determined fiber architecture within a hydrogel. Bas et al. combined the ECM-like structure of hydrogels with highly oriented PCL fibers fabricated by Melt Electrospinning Writing (MEW). Stacked fibers with lay- down patterns were prepared by MEW, and then infiltrated with hydrogels in custom-made molds and crosslinked by a reduction–oxidation initiating system. The fiber reinforced hydrogels achieved an outstanding increase in the mechanical properties at high strain rates.

The fiber reinforced hydrogels achieved an outstanding increase in the mechanical properties at high strain rates. However, the thickness of the scaffold was limited to 1 mm.This paper presents a novel divergence electrospinning strategy to rapid assemble aligned nanofibers along Z-axis by two axisymmetric and separately grounded bevels. This phenomenon was attributed to the whipping of the continuous electrospun jet in a symmetrically diverged electric field. While most of the nanofibers attached on either bevel of the collector, some remained between the two bevels due to the dragging effect of the whipping. As the fiber density increased, the electrostatic repulsive forces resulted from the residual charges on the fibers enhanced, driving the2401subsequent electrospun fibers away from the existing bundles, resulting in a fast build-up along the z-axis.The nanofibers were uniaxial aligned in the 3D scaffold.

The results from fiber density analysis revealed a gradient of fiber density from top to bottom layers: a decrease of fiber density from top to bottom, as well as from periphery to the center. We hypothesized that this phenomenon was due to the repulsive effects attributed to the residual charges on the electrospun nanofibers. For regular electrospinning, the nanofibers are deposited on a grounded surface, where the repulsive forces (FR) from the accumulated residual charges are fractional compared to the static electric force (FE) until the fiber density and thickness reach a very high level (Figure 8a). In our case, the bridging nanofibers were hanged between two grounded bevels. The static electric forces exerted on the nanofiber were diverged to two directions, led to a weaker resultant force downward (Figure 8b).

The static electric forces exerted on the nanofiber were diverged to two directions, led to a weaker resultant force downward (Figure 8b). When nanofibers accumulated, the repulsive forces between the fibers might soon exceed the static electric forces and thus pushed the later electrospun fibers to the peripheral space. Once the nanofibers at the peripheral areas became denser, it would be more difficult for the electrospun jet reach the central area. Therefore the gradient of fiber density would only increase as electrospinning continued.Figure 8. Force analysis of a single electrospun fiber in (a). a regular electric field, and (b). a divergent electricfield. (c). A simulation of the fiber density along the z-axis of the collector bevel.Cells were attached on the microfibers and grew along the fibers. This result is consistent with many previous studies on aligned fiber scaffold.

This result is consistent with many previous studies on aligned fiber scaffold. It was reported that aligned micro- and nanofibers are favorable in replicating the ECM for specific tissues which is composed of perpendicularly interwoven collagen strips. An aligned electrospun fibrous scaffold can provide contact guidance to migration and extension, resulting in an elongation and alignment of cytoskeleton and nuclei along the axes of the fibers [8]. The effect of contact guidance of aligned electrospun fibers on cell morphological changes was evident in other many cell types [9, 10]. The scaffold developed in this study can be directly embedded in cell-laden hydrogel.

The scaffold developed in this study can be directly embedded in cell-laden hydrogel. This technique can be applied in engineering of musculoskeletal soft tissues such as tendons, ligaments, knee menisci, etc.[11], where fibrous cytoskeletal organization is critical for tissue formation and functions.Future work will focus on how does the electrical and rheological properties of polymer solution determine the microfiber resolution during the dynamic stretching induced by the divergence2402electric field. Most of the existing models for electrospinning focus on the rudimental configuration in which the microfibers are randomly electrospun onto a grounded surface driven by an ideal point- to-plane electric field.

It is critical to establish the relationships between material properties, the finite extensibility of the polymer chains, and the microfiber diameter.ConclusionThis paper presented a novel divergence electrospinning strategy with double-bevel collectors for 3D nanofiber scaffold fabrication. A highly aligned and porous nanofiber scaffold with a size in centimeter scale was rapidly fabricated through this technique. A porosity (fiber density) gradient was observed along the z-axis of the scaffold. The aligned fibrous structure induced an aligned cell attachment and growth. This study provided an innovative rapid fabrication approach for biomimetic microstructures and showed a great potential as a future additive microfabrication platform for tissue engineering.24031.2. 3.4.5. 6.7.8.9.10.11.ReferencesMichel, G., et al., The cell wall polysaccharide metabolism of the brown alga Ectocarpus siliculosus.

brication of bone substitute material by Rapid PrototypingA. Ott, J. Heinzl, D. Janitza, R. Pelzer Lehrstuhl für Feingerätebau und Mikrotechnik, Technische Universität MünchenAbstract:Bone tissue engineering has gained much attention in recent years. A key requirement in this field is the development of scaffold structures, on which cells adhere. This can be done by fabricating scaffolds by direct procedures like 3D-printing or by indirect procedures like casting. With the 3D-printing process different structures were build up by using hydroxyapatite powder (HA) and a special binder material. Afterwards the printed 3D structures were sintered. For the casting process molds have been made of different resins by stereolithography and other processes using polymers and waxes. These structures were filled by a suspension of HA. By heating the resulting polymer/ceramic composite to a specific temperature it is possible to combust the polymer or wax.

These structures were filled by a suspension of HA. By heating the resulting polymer/ceramic composite to a specific temperature it is possible to combust the polymer or wax. By further heating the remaining body, the HA is sintered. Compared to the 3D printing a better resolution can be obtained here. But there are restrictions regarding the ratio of polymer and the HA ceramic during the heating process which means a limitation for the level of porosity.IntroductionIn the field of biomedical implants attempts to design suitable systems have been made from different directions. One current research field focuses on tissue engineering. The research group Fortepro is developing a processes for fabricating bone substitute material by using rapid the prototypingtechniques.

One current research field focuses on tissue engineering. The research group Fortepro is developing a processes for fabricating bone substitute material by using rapid the prototypingtechniques. Theseimplants are designedto replace bone defects atData Acquisition1Data preparation2Source : Howmedica Leibinger GmbH & Co.KG Source : Howmedica Leibinger GmbH & Co.KG3Implant fabrication Process of Custom Implants by Rapid Manufacturing4Settling with cellsImplantFigure 1. Principle process sequence7845head and musculoskeletal system. They are made of hydroxyapatite (HA) and will be replaced by endogenous bone material after implantation by building new tissue. Figure 1 shows the complete process sequence. In a first step data from the defect is collected with computer tomography (CT). Then the three-dimensional geometry of the implant is designed with a special software and converted in a standard triangulated language (STL) file.

Then the three-dimensional geometry of the implant is designed with a special software and converted in a standard triangulated language (STL) file. With different RP processes the implant geometry is produced and later prepared with cells. Afterwards the prepared scaffold can be implanted.A key requirement is the development of scaffold structures on which cells can adhere. The aim of the institute “Feingerätebau und Mikrotechnik” is to fabricate these scaffolds with controlled internal porosity.Requirements on the implantsThis study aims at the creation of a porous artificial extracellular matrix or scaffold to accommodate cells and guide their growth and tissue regeneration in three-dimensions (3D). These scaffolds should be created individually for each patient, with interconnected and controlled porosity and pore distribution.

These scaffolds should be created individually for each patient, with interconnected and controlled porosity and pore distribution. Furthermore the mechanical strength not only depends on the 3D structure of these scaffolds but also on the used process parameters for example the temperature profile during the sintering process. An overview of the requirements on the 3-D structures are given in Figure 2.Requirements on the 3D-structure: – Component accuracy – Defined inner structure for cell ingrowthMechanical requirements: – Elastic modulus – Breaking strength – Fatigue strengthProcessMaterialRapid Prototyping (RP) : - 3D -Printing (3DP) - CastingProcess requirements:– Powder form and particle size – Grain size distribution – Binder for volumetric micro dosing – Slurry – Polymer mouldMedical requirements: – Biocompatibility – Degradation – ProliferationFigure 2. Requirements of the processPart accuracy:The process allows to generate patient individual implants.

Requirements of the processPart accuracy:The process allows to generate patient individual implants. There is no need of post processing the implant by the medicines at the later surgery. Moreover, it is necessary for cell growth and flow transport of nutrients and metabolic waste that the defined inner structure and porosity can be produced consistent. The desired parts should have interconnecting pores with a size of about 400 to 600 µm for cell proliferation and the porosity should be around 50%.785Mechanical requirements:Desired is an elastic modulus, breaking strength and fatigue strength with valves similar to those of natural bones. Our research group wants to develop an implant which can be handled by the medicines without exhaustive care and will serve for implantation in regions without high loading eg. the head region. Medical requirements:Biocompatibility is necessary for the implantation.

the head region. Medical requirements:Biocompatibility is necessary for the implantation. Furthermore the scaffold structures should allow the infiltration of stem cells. This is necessary to replace the biodegradable artificial bone substitute material by autogenous material. Literature shows that the hydoxyapatite ceramics have been well established because of their high biocompatibility qualities. The degradation of the implant also depends on the used material and is so far not clarified completely. Process requirements:For the different processes the used materials have big influence on the results. For the 3D printing process the particle size and a spherical particle form are very important. Furthermore it is necessary to prevent the particles from agglomeration. A lot of printing experiments were done to investigate the influence of different HA and binder combinations.

Furthermore it is necessary to prevent the particles from agglomeration. A lot of printing experiments were done to investigate the influence of different HA and binder combinations. On the other hand the quality of the filling of the negative molds depends strongly on the prepared slurry. Different proportion of HA ceramic, deflocculant, solvents and water have big influence on the results. Moreover the slurry and the temperature profile can be matched to the different materials of the molds.Processes 3D-printingThis process creates parts by a layered printing process with adhesive bonding, using HA powder as a base material. After loading the STL-file, which defines the geometry and the interconnecting pores, the software of the Rapid Prototyping machine slices the three- dimensional data into two-dimensional pictures according to the cross-sectional area of the object. Each layer of powder is selectively joined where the scaffold is to be formed by ink-jet printing.

Each layer of powder is selectively joined where the scaffold is to be formed by ink-jet printing. This process is repeated layer by layer until the green body of the scaffold is complete. Afterwards the powder which is not joined together is removed and the green body is sintered.This printing process requires a special powder. The resolution depends on the form and particle size and the grain size distribution. The major problem is to choose an appropriate binder for micro dosing. The used hydroxyapatite powder is spray dried with particle sizes between 100–200µm diameter. The binder, which joins the powder particles, was also specially developed for this process. Each powder layer is about 200µm thick and finished with a counter rotating roll, which sleeks the HA. The micro dosing of the binder is realized with a piezoelectric drop on demand system or an electromagnetic valve.

The micro dosing of the binder is realized with a piezoelectric drop on demand system or an electromagnetic valve. Therewith it became possible to dose drops with a diameter of 60-80µm.The temperature profile for sintering only depends on the ceramic. Starting at room temperature the temperature is increased to 1250 °C at 100 °C/hour. This temperature is hold for about 1 hour min and then the ceramic is cooled down to room temperature with about 100 °C/hour.786Results 3D-printingThe first printed bodies have been fabricated without internal structure. These tests proved, that the green bodies can be sintered and led to an overview of the quality of the printing results by using different powders and binders. On scanning electron microscope pictures, (Figure 3) the sinter-necks between the agglomerates can be observed as well as the sintered powder-particles in the agglomerate.

On scanning electron microscope pictures, (Figure 3) the sinter-necks between the agglomerates can be observed as well as the sintered powder-particles in the agglomerate. These results prove the feasibility of the process.Sintered powder particlesSinter-neckFigure 3. Sintered necks between the spray dried particlesFigure 4 (left) shows a printing result of a binder/ceramic combination where the printed plates were bended up at the edges and the edges themselves were damaged. The right side of this figure shows the results of a very good ceramic binder combination but with bad process parameters (speed, droplet size, amount of binder etc.).Figure 4. Printed HA ceramic plates with different process parametersThe next step was the production of scaffolds with defined inner structures in three dimensions. These parts are demonstrated in Figure 5.

Printed HA ceramic plates with different process parametersThe next step was the production of scaffolds with defined inner structures in three dimensions. These parts are demonstrated in Figure 5. The size is 10mm x 10mm x 6mm which fits perfectly in cell seeding equipment and is made of rectangular beams with a cross section of about 0.5mm x 0.5mm. The strength of the printed parts (green bodies) is low but after sintering at 1250°C they are solid enough for handling. The powder-binder interaction influences the strength of the part and the behaviour of the wetted powder. The theoretical achievable strength and the strength of the printed scaffolds is still under investigation.7871 cmFigure 5. 3D-printed ceramic parts with internal porosityhe scanning electron microscope pictures (Figure 6) shows the cross section of a printed and T sintered scaffold. The surface remains very rough and spherical ceramic particles are visible.

The surface remains very rough and spherical ceramic particles are visible. At the fractured surface of the ceramic micro porosity is remaining.Figure 6. Micro porosity at the fractured surface of a printed HA-ceramic788Processes CastingAfter presenting the results of the 3D printing process now the second RP process is described. Figure 7 shows the flow chart of the implant manufacturing process by casting. On the left side the process of the hydroxyapatite suspension preparation can be seen. This is done by the Friedrich Bauer Institute which is highly experienced in ceramic materials.Mold Designing from Negative Implant ImagesProducing of HA - SlurryFabrication of Molds by RP-ProcessesImpregnationHeating, Mold and Binder RemovalSinteringFigure 7. Flow chart of the manufacturing processIn this chapter the designing of molds from negative images of implants and the fabrication of the molds by RP processes is described.

Flow chart of the manufacturing processIn this chapter the designing of molds from negative images of implants and the fabrication of the molds by RP processes is described. At the beginning of the project we focused on the stereolithography as the highest accuracy for getting the desired internal structure of the molds were expected. But also, the 3D Wax printing process is under investigation. In the next step the prepared slurry is filled into the mold (Impregnation of the mold). Followed by a thermal process to remove the binder components from the suspension and to combust the mold. By rising the temperature up to 1250° Celsius the HA ceramic is sintered. In this cage the sinter profile has to be adjusted not only to the ceramic but also to the material of the mold. The heating for combusting the molds has to be done very slowly with holding the temperature at a special point.

The heating for combusting the molds has to be done very slowly with holding the temperature at a special point. After the combustion of the mold is completed, the heating process for sintering is comparable with the temperature profile described for the 3D printing process. The critical points are the correct impregnation (filling of the molds) and the prevention of cracks in the ceramic during the first heating process, when the mold causes stress because of its thermal expansion.Figure 8 demonstrates the feasibility of this process. The mold was fabricated by stereolithography and filled by a standard procedure of slip casting. The geometry of the parts is similar to the geometry used in the 3D printing process. The rough surface due to the layer by layer fabrication on the mold (left) can also be observed at the sintered scaffold (right).

The rough surface due to the layer by layer fabrication on the mold (left) can also be observed at the sintered scaffold (right). But the cracks in the structure led to a failure of the whole scaffold.789Figure 8: ESEM micrographs of the RP mold and the resulting scaffoldThe interconnecting pores size for cell proliferation and the maximum wall thickness for degradation are limiting the design of the implants. To avoid cracks due to thermal stress, the ratio between resin and ceramic had to be decreased. This was realized by inserting hollow structures in the mold which are not filled by the ceramic slurry. This way, the stress on the ceramic during combusting the molds can be reduced. But still, there was a high reject due to defects. Therefore a new design of the parts and a better filling process were established.Filling processFigure 9 shows a micro computed tomography (µ-CT) of an impregnated mold prepared at the beginning of the project.

The structure in dark-grey is the shot of the mold. The white color presents the ceramic. The origin of the cracks during the drying of the ceramic can be traced very well. The insufficient filling of the mold made it necessary to improve this process. By adding different tensides to the slurry and by investigating different filling methods the characteristics of the slurry have been changed.Figure 9. Mikro CT of an impregnated mold790The vacuum-, the centrifuge- and the pressure/vacuum methods have been investigated. The pressure/vacuum method (see Figure 10) solved most of the problems during the impregnation of the molds. A small chamber is filled with the HA slurry and the mold. By exceeding pressure on one side and low pressure on the other side of the chamber the small cavities could be reasonable filled without any air bubbles. By using low pressure it became possible to dry the slurry during the filling process.

By using low pressure it became possible to dry the slurry during the filling process. This way the swelling of the resin molds by the water-based slurry can be reduced.Figure 10. The exceed pressure / vacuum filling stationThe used materials for the molds are standard resins for stereolithography and standard wax of 3D printing respectively. Therefore one major question is the influence of the toxic vapor on the ceramic parts during combustion the mold material. This is done by technical analysis like EDX or by analyzing the cell proliferation on the material. EDX analysis of sintered ceramics produced with different stereolithography materials and waxes have been curried out (Figure 11).Figure 11. EDX spectrum of a molded ceramic scaffold791No toxic residua could be found at the surface of the ceramic parts. First experiments with cell p roliferation on the surface of the sintered parts show good results.

First experiments with cell p roliferation on the surface of the sintered parts show good results. Next investigations with stem cells have to prove if the chosen wall and pore size are adequate for cell proliferation on the sintered implants.esults of the casting process The desired channel and wal l size of less than 500µm can easily be achieved by using the SLS te chnology. But due to the thermal expansion and the exited stress it is necessary to keep a ratio of polymer to ceramic of less than 50%. Using a honeycomb structure with hollow blocks, allows the reduction of the polymer ratio dramatically. Therefore the filled mold is more suitable for theRFigure 12. Ceramic parts with internal porositythermal process. Moreover sharp edges can be reduced which is expected to be better for cell roliferation and the filling process. p In contrast with the rough surface of the 3D printing scaffolds the results by casting are very smooth.

p In contrast with the rough surface of the 3D printing scaffolds the results by casting are very smooth. Figure 13 shows a ESEM m icrograph of the scaffold with a very dense structure and nearly no micro-porosity. On the left side, the single layers of the molds are also visible in the scaffold.Figure 13. High-magnification SEM-micrograph of scaffold surface792ConclusionBoth RP processes are suitable for producing scaffolds for tissue engineering. First investigations proofed the possibility of cell proliferation on the surface of the structures. Further steps will include the habitation of the scaffold structures with stem cells and investigate the degradation of the ceramic and bone morphogenic protein induced osteogenesis.Acknowledgement:We thank the “Bayrische Forschungsstiftung” for founding the project Fortepro.

BRICATION OF CELLULAR CORDIERITE PREFORMS VIA BINDER JETTINGDean Snelling1, Christopher Williams1, Carlos Suchicital2, Alan Druschitz2 1Design, Research, and Education for Additive Manufacturing Systems Laboratory; Department of Mechanical Engineering 2Department of Materials Science and EngineeringVirginia Tech, Blacksburg, VAABSTRACTMetal Matrix Composites (MMCs) combine two dissimilar materials – a metal as the matrix and a ceramic as the charge or insert to provide unique properties, e.g. low density, high specific strength, high specific modulus, and wear resistance. The ceramic inserts in these composite materials are limited to ceramic fibers and open cell stochastic ceramic foams due to geometric constraints imposed by traditional manufacturing processes.

The ceramic inserts in these composite materials are limited to ceramic fibers and open cell stochastic ceramic foams due to geometric constraints imposed by traditional manufacturing processes. The geometric design freedom offered by Additive Manufacturing (AM) could enable a designer to realize ceramic preforms with complex cellular geometries that are designed to achieve multiple functions (e.g., low mass and increased stiffness). The goal of this work is to explore the use of Binder Jetting as a means of fabricating cordierite parts of designed mesostructure for use as ceramic preforms for MMCs. In this paper, the authors describe their exploration of the appropriate printing process parameters and post-process sintering parameters that enable successful fabrication of complex cordierite artifacts. Measurements of bulk density, linear shrinkage, porosity, and x-ray diffraction are conducted on pre- and post-sintered printed cordierite structures.1.

Measurements of bulk density, linear shrinkage, porosity, and x-ray diffraction are conducted on pre- and post-sintered printed cordierite structures.1. IntroductionMetal matrix composites are defined as a metal matrix reinforced by a constituent material, usually a ceramic [1], to provide materials with improved properties including low density, high specific strength, high specific modulus, high thermal conductivity, and wear resistance [2]. The matrix of the composite supports and transmits loads distributed to the secondary material used as reinforcement [3]. Metal matrix composites (MMC) can be broken down into two different fabrication processes: 1) solid state processing and 2) liquid state processing. Solid state processing includes using powder metallurgy and diffusion bonding. Liquid-state processing involves infiltrating, dispersion, spraying, and in-situ fabrication [2].

Solid state processing includes using powder metallurgy and diffusion bonding. Liquid-state processing involves infiltrating, dispersion, spraying, and in-situ fabrication [2]. Many of the applications where metal matrix composites could have a significant impact have not been adopted because of the high cost to performance ratio [4] associated with the ceramic insert (e.g., fibers) used to create the reinforcement. For the case of ceramic fibers, the preforms are manufactured by mixing the fibers by press forming or suction forming. The liquid is removed and the preform is dried [5]. Such manufacturing processes are complex and the resulting parts are not reliable and can have unwanted phases. Furthermore, open cell ceramics (ceramic foams) have been used for the formation of MMCs to decrease the cost associated with preforms made with ceramic fibers [4].

Furthermore, open cell ceramics (ceramic foams) have been used for the formation of MMCs to decrease the cost associated with preforms made with ceramic fibers [4]. But like MMCs fabricated with ceramic fibers, cost and geometric limitations are imposed by the manufacturing process associated with machining the final parts. To alleviate these limitations, the authors propose the use of Binder Jetting additive manufacturing technology (AM) to manufacture complex ceramic cellular preforms for thecreation of MMCs. Using AM to form the ceramic pre-sintered pieces for MMCs would enable a manufacturer the freedom in complexity to create unique composites designed for specific applications.

Using AM to form the ceramic pre-sintered pieces for MMCs would enable a manufacturer the freedom in complexity to create unique composites designed for specific applications. Traditional techniques for forming ceramics (pressing, casting, plastically forming the ceramics into a green part, followed by sintering in a pressure-less or pressure-assisted cycle [6]) are unable to manufacture complex geometries such as ordered cellular geometries (repeating unit cells featuring voids ordered throughout the structure).

Manufacturing ordered cellular preforms for MMCs via AM would provide multiple functionality to the resultant part including decreased mass while increasing strength and stiffness and would allow an engineer to produce intricate geometries for specific design purposes (as opposed to that found in stochastic ceramic fiber and open cell preform techniques).1.1.Binder Jetting Process and CapabilitiesBinder Jetting is an AM technology that creates artifacts through the deposition of binder into a powder bed of raw material (Figure 1). Once a layer has been printed, the powder feed piston raises, the build piston lowers, and a counter-rotating roller spreads a new layer of powder on top of the previous layer. The subsequent layer is then printed and is adhered to the previous layer by the jetted binder. Binder Jetting is a versatile method of printing powdered materials.

The subsequent layer is then printed and is adhered to the previous layer by the jetted binder. Binder Jetting is a versatile method of printing powdered materials. Any powder material that can be synthesized, deposited, and bound with a jettable binder may be used [7] including plaster, sand, metal, and other ceramics [8–10]. When printing ceramic powder, the resultant green part is sintered (and the binder pyrolyzed) to realize the final ceramic artifact.Figure 1. Binder Jetting AM SchematicOther AM technologies such as powder bed fusion [11], vat photo polymerization [12], and extrusion [13] have been used to process ceramics, but they are outside the focus of this work. Binder Jetting is a focus of this work as it offers a low-cost and scalable means of producing ceramic artifacts with minimal constraints imposed on part topology.

Binder Jetting is a focus of this work as it offers a low-cost and scalable means of producing ceramic artifacts with minimal constraints imposed on part topology. The technology is commercially offered by ExOne’s M-Print® and M-Flex® machines [14] and 3D Systems’ Professional ProJet® [15].Ceramic parts have been successfully created via Binder Jetting to meet a variety of needs. Yoo and coauthors (1993) printed alumina powders to manufacture structural ceramic components that were 99.2% dense and had a flexural strength of 324 MPa [10]. Additionally, they were able to control the microstructure of zirconia toughened alumina (ZTA) by adding a stabilizer during the jetting process [16]. Seitz and coauthors’ (2005) used Binder Jetting to print hydroxyapatite (HA) ceramics to produce ceramic scaffolds for bone tissue engineering [17].They were able to attain internal passage resolutions of 450 µm, wall thicknesses of 330 µm, and 22 MPa in compression.

Sachs and coauthors (1990) tested the dimensional accuracy and part strength of printed parts from several ceramic powders including 320 grit (45 μm) aluminum oxide, silica, zirconia, zircon, and silicon carbide. Dimensional control was determined to be approximately ±20 μm along 38 mm dimensions. Part strength was determined to be strong enough to handle and survive during pouring, but not strong enough to overcome the contraction of the metal during cooling. They determined that the geometric freedom offered by Binder Jetting could be used to efficiently prototype materials with specific application considerations and material properties [18].Furthermore, AM of ceramics has been used to print preforms, but not specifically ordered cellular preforms. Sachs and coauthors (1990) explored the ability of Binder Jetting to create green ceramic preforms for infiltrating molten metal in order to manufacture metal matrix composites [18].

Sachs and coauthors (1990) explored the ability of Binder Jetting to create green ceramic preforms for infiltrating molten metal in order to manufacture metal matrix composites [18]. Additionally, they printed macroscopically toughened composites from SiC and Al 224 to eliminate cracks from propagating due to the interface [19]. Moon et al. (2001) printed carbon preforms with a designed mesostructure for reaction-infiltrating creating SiC-Si composites [20]. Unlike Sachs and coauthors, they utilized two materials to create a chemically bonded MMC. The print head speed was varied to obtain the desired preform microstructure and was controlled along the length of the complex preforms therefore controlling the carbon content along the part. This lead to the ability to vary the reaction along the part and produce functionally graded SiC-C composites. The part contained three times the amount of binder saturation from one end to the other, resulting in a Si content of ~8 wt.

The part contained three times the amount of binder saturation from one end to the other, resulting in a Si content of ~8 wt. % at the end with the highest binder saturation compared to ~ 24 wt. % at the end with the least binder. Unlike the work by Sachs and coauthors, and Moon et. al., this work uses precursor materials to generate cellular shapes for ceramic inserts.1.2.Overview and Purpose of WorkThe goal of this work is to form (print), sinter, and characterize ceramic cellular parts as precursors for MMCs. Specifically, an ExOne R2® system was used to fabricate cordierite ordered cellular structures. By using cellular geometries as MMC preforms, the authors strive to provide sufficient strength in bending and compression modes while decreasing the weight of the part [21]. A schematic of the manufacturing process is shown in Figure 2.Figure 2.

A schematic of the manufacturing process is shown in Figure 2.Figure 2. Proposed ceramic preform fabrication processAn overview of the materials and experimental methods employed in this study is presented in Section 2. The results from raw material characterization, printing parameters for the green part manufacture, sintering experimentation, and post-sintered characterization are given in Section 3 Closure is offered in Section 4 along with a discussion of future considerations.2. Material Selection and Experimental Methods2.1.Powder SelectionIn order to manufacture suitable ceramic inserts for fabricating MMCs, the precursor materials were first selected. The materials were chosen based on the functional requirements of the manufacturing process (e.g., printing, sintering, and casting requirements).

The materials were chosen based on the functional requirements of the manufacturing process (e.g., printing, sintering, and casting requirements). Several candidate materials were considered after successfully printing with the ExOne R2® system, yet, these did not meet the functional requirements of the manufacturing process. Since the MMCs are fabricated via metal casting, the final ceramic pieces were required to have high thermal shock resistance to eliminate the possibility of cracking during manufacture. Additionally, the parts needed to have lower sintering temperature to match the facility resources readily available at Virginia Tech.Due to its poor thermal shock resistance behavior, Al2O3 alone was not considered a useful candidate for higher temperature cast metal matrix. However, when reacted with silica (SiO2) and magnesia (MgO), the resultant material, cordierite (2MgO•2Al2O3•5SiO2), can provide amuch improved thermal shock resistance.

However, when reacted with silica (SiO2) and magnesia (MgO), the resultant material, cordierite (2MgO•2Al2O3•5SiO2), can provide amuch improved thermal shock resistance. Cordierite was chosen because it has excellent thermal shock properties to withstand molten metal temperatures and also has a sintering temperature in the range of the electric furnace available [22]. In addition, the presence of a liquid phase can encourage higher final densities and decrease the required sintering temperature. Additionally, varying combinations of these materials would provide different properties for discrete applications and could be easily adjusted to fit further applications. For example, if increased strength is required and a sintering temperature is not a concern, the composition can be varied based on the properties of the resulting material.

For example, if increased strength is required and a sintering temperature is not a concern, the composition can be varied based on the properties of the resulting material. For these reasons cordierite was chosen as the candidate material.2.2.Powder PreparationCordierite precursor materials were mixed using the amounts in Table 1 (derived from the molar composition given in Section 2.1). The raw materials were mixed in a container with Al2O3 pellets to ensure thorough blending and removal of agglomerates.Table 1. Ceramic raw material mixturesCeramic Cordierite (2MgO ·2 Al2O3 · 5 SiO2)MgO (wt. %) 13.78Al2O3 (wt. %) 34.86SiO2 (wt. %) 51.36The minimum mean particulate powder size for Binder Jetting is approximately 20 μm [23] and ideal shape for powder flow is spherical. However, off-the-shelf oxide powders for creating sintered ceramics are mostly angular containing smaller mean size distributions.

However, off-the-shelf oxide powders for creating sintered ceramics are mostly angular containing smaller mean size distributions. Precursor powders typically are pressed and sintered to increase the final density with less shrinkage of the final part. A smaller mean particle size aids in reactivity, grain size, density requirements, rate of sintering, and temperature [24]. Printing ceramics via Binder Jetting is limited by the ability of the powder to be recoated by the binder evenly across the build tray. Dry powders with small particle sizes (<20 μm) are difficult to spread due to the cohesive strength of the high surface area of fine particles [10,23]. Additionally, spherical powders are desired for ease of spreading as opposed to angular particle shapes [19].For this work, off-the-shelf ceramic powders were used as the raw materials for printing in the ExOne R2® system.

To ensure their compatibility with the system, the particle size distribution of the precursor materials and the precursor mixture were analyzed using a Horiba LA-950® Laser Scattering Particle Size Distribution Analyzer and reported in Section 3.1.2.3.Binder Jetting Printing ParametersThe goal of the proposed forming process was to manufacture geometrically intricate and accurate parts with sufficient green part strength for further handling, e.g., cleaning and sintering. During the forming process, printing parameters such as binder selection, layer thickness, heater temperature, binder saturation, spread speed, and drying rate were adjusted to achieve these objectives. These settings were determined by experimentally printing and examining green parts post-printing.

These settings were determined by experimentally printing and examining green parts post-printing. For example, small rectangular parts were printed to ensure that the green parts had sufficient part strength to cure and be transferred to the furnace for sintering.Additional octet truss unit cells (section A-B in Figure 3) were printed to determine the printing parameters of more complex shapes.Section A–BFigure 3. CAD Schematic of cellular repeating octet (A-B dotted line) truss pattern.The printing parameters were adjusted to produce two distinct part types; i) solid parts for characterization purposes and ii) complex cellular parts for the purpose of creating metal-ceramic cellular structures (Section 3). It was determined by experimentally printing these two types of shapes that the binder saturation was different. Binder saturation was adjusted based on resultant green part strength.

It was determined by experimentally printing these two types of shapes that the binder saturation was different. Binder saturation was adjusted based on resultant green part strength. Excessive binder would decrease the quality of the final parts due to unwanted binder bleeding, while green parts with too little binder could have limited strength.Furthermore, the drying rate produced by the overhead heater on the printer, was dependent on binder saturation values. The drying rate was the rate the overhead heater moves across the build tray after each layer is printed. The overhead heater on the AM system was used to dry binder after each layer to avoid excessive bleeding of binder by partially curing the binder.The layer thickness was determined experimentally after printing the materials at different layer thicknesses. Any spreading defects during printing of thinner layers would produce potentially catastrophic results in the previous layer and therefore in the final part.

Any spreading defects during printing of thinner layers would produce potentially catastrophic results in the previous layer and therefore in the final part. In addition, the larger layer thickness provided much shorter overall build times. Because of the limitations of spreading fine angular powders (Section 3.1) thicker layers (200 µm) were tried and found to yield higher quality parts.As previously noted, the binder used was ExOne’s standard thermosetting binder that cures at 190°C. Once printing was completed, the green part was removed from the build bed, depowdered, and placed in an oven at the curing temperature for 2 hours. The cured binder gave strength to the part before placing in a high temperature environment for sintering.2.4.Part SinteringSmall rectangular green parts that were used for testing printing parameters were also used for several firing trials to determine a satisfactory sintering temperature.

Ceramic green specimens were placed in a furnace for firing with a 5°C/min ramp rate to the binder burnout temperature of 800°C for one hour, a 5°C/min ramp rate to the final sintering temperature with a dwell of one hour. Longer sintering cycles were tested and it was determined that change in temperature has a significantly larger effect on diffusion. The fired samples were removed from the furnace, at temperature (1350°C, 1360°C, 1375°C, and 1400°C) and at the end of the dwell period. The pieces were fired until parts acquired a glassy texture or appeared to be fully melted.This observation determined a rudimentary range of temperatures for finding the preferred firing cycle.

The pieces were fired until parts acquired a glassy texture or appeared to be fully melted.This observation determined a rudimentary range of temperatures for finding the preferred firing cycle. Relative bulk density and shrinkage characterization (Section 3.3.2) was performed on specimens sintered at that range of temperatures.2.5.Ceramic CharacterizationIn addition to relative bulk density and shrinkage, the sintered ceramic parts were submitted to x-ray diffraction (XRD) and Optical Microscopy in order to determine the properties of the fired ceramic. The relative bulk density of the sintered ceramic parts was evaluated using the Archimedes Principle. Because the density determines the ultimate strength of the ceramic, this metric (along with shrinkage and porosity) was used to determine the sintering cycle.

Because the density determines the ultimate strength of the ceramic, this metric (along with shrinkage and porosity) was used to determine the sintering cycle. The parts were impregnated with oil before testing to seal the surface porosity and eliminate any water from penetrating interconnected porosity of the part (ASTM standard B962) [25].The physical dimensions of the green parts were compared to those from the sintered part to obtain the apparent shrinkage. Additionally, the measurements from the CAD model were compared to those of the green and sintered parts to obtain the correct final part when modeling in CAD.Optical microscopy was used to confirm the densification of the ceramic structure with each sintering cycle. Each printed ceramic block was mounted in an epoxy media and ground to 600 grit SiC paper, cleaned and examined.

Each printed ceramic block was mounted in an epoxy media and ground to 600 grit SiC paper, cleaned and examined. Optical micrographs of representative samples are shown in Section 3.2.XRD analysis was completed on the final parts to determine how close the sintered ceramic sample’s crystallographic structure came to the data found in the literature for a cordierite ceramic material. All samples to be examined were first ground to powder form with a mortar and pestle.3. Results3.1.Powder CharacterizationThe precursor powders were mixed and first tested with ExOne’s standard binder (PM-B- SR2-05) used in the ExOne Binder Jetting AM equipment, to ensure that the material and the binder were compatible.

The binder and precursor materials were mixed by hand, pressed into pellets, and sintered to verify that the parts, when printed, would retain their shape after sintering.Following this verification of the binder/powder compatibility, the powder size distribution was analyzed to ensure that the powder could be successfully recoated during printing. The mean particle sizes of the precursor powders are shown in Table 2 and the number fraction (bar chart) and cumulative percent finer distributions (red line) of the mixed material are shown in Figure 4.7MgO & SiO2120)%( n o i t c a r F r e b m u N654321Mean Size = 30Al2O310080604020r e n i F t n e c r e P r e b m u N e v i t a u m u Cl00Particle Diameter (µm)Figure 4. Particle size analysis: Number fraction and cumulative number percent finer distributions of cordierite precursor raw materials.Table 2.

Particle size analysis: Number fraction and cumulative number percent finer distributions of cordierite precursor raw materials.Table 2. Mean particle sizes of individual materials and cordierite precursor mixtureMaterial Mean Diameter (µm)Magnesia (MgO) Alumina (Al2O3) Silica (SiO2) Precursor Mixture11.0667.8916.8030.47The cordierite precursor powder mixture had a mean particle size (30.5 µm), which is considered satisfactory for spreading in the AM system. However, not all individual materials were above the minimum size threshold (20 µm) (Table 2) and the trimodal distribution of particles in combination with the angular shape of the raw materials, resulted in a limited homogeneity of the powder mixture spreading step.

High binder saturation values were not appropriate for solid objects, due to binder bleeding, which would decrease dimensional accuracy by excess bound powder from the surrounding support powder. Binder saturation was based on a printer input parameter – initial packing density. The ExOne R2® system was limited to the amount of additional pressure provided by the recoater to create denser “as printed” parts. Typically, spherical powders were used for better spread ability (Section 2.1) and have similar tap and apparent densities. The tap density could be measured and input into the printer. Additionally, adjusting the particle size distributions (mono-sized particles, bimodal, ternary, quaternary etc.) could be used to increase the initial packing density of the printed material.Heating the jetted binder gave the printed part strength during and after printing the part in its green (unfired) state.

The larger the saturation values (and consequently more binder) required more heat for curing. Solid parts required slower drying rate for the large volume of material to be dried/cured while complex parts used an accelerated rate to avoid over-drying and delamination of layers during depowdering the green part.As discussed previously, spreading was satisfactory and was minimally controlled by adjusting spread speed. Setting the spread speed to the slowest value of 1 mm/s provided the best spread of the precursor materials. A slow spread speed minimized agglomerates from being drug across the new layer and instead created a tumble motion from the wave of new powder.

A slow spread speed minimized agglomerates from being drug across the new layer and instead created a tumble motion from the wave of new powder. After experimentation, the powder was able to spread evenly at low spread roller speeds (Section 2.3) with minor tracks from the previous due to agglomerates forming from the small particles in the wave of powder forming the new layer.3.3.Sintered Part Characterization3.3.1.Relative bulk density measurementsBased on the raw materials utilized in the precursor powder mixture for the AM process, the resulting fired ceramic part was expected to show a cordierite or cordierite like nature. Ceramics were found to have a glassy surface at approximately 1400°C but was too weak at 1350°C, therefore a temperature range of 1350°C-1400°C for sintering trials as discussed in 2.4.Figure 5 shows the relative bulk density measured at different sintering temperatures and is based on a theoretical density for cordierite (2.51 g/cc)[26].

As expected, the density increased significantly until reaching approximately 92% at 1375°C. At this temperature, the ceramic parts reached the final stage of sintering (~90% density) and any further increase in temperature would generate an excess of the bonding liquid phase (with lower viscosity) and the integrity of the part would be compromised by slumping or by a complete melting [6].Figure 5. Relative bulk densities of fired ceramic parts as a function of the firing temperature3.3.2.Shrinkage measurementsThe shrinkage for each dimension as a result of sintering is shown in Figure 6. It was expected, given the correlation with the bulk density, that the shrinkage would be significantly greater at higher sintering temperatures. But it was also noted that with large shrinkage values, the ceramic parts experienced significant warping for both solid and complex objects. In both cases, the parts had a larger shrinkage in the Z-direction as compared to the X and Y directions.

In both cases, the parts had a larger shrinkage in the Z-direction as compared to the X and Y directions. This is due to the effects of gravity on the part during the sintering process and can be corrected to meet geometrical dimensions by compensating the green part.Figure 6. Ceramic shrinkage for ceramic parts at various sintering temperatures3.3.3.Optical microscopy of printed ceramicsA macro sectioned piece and optical micrographs of representative ceramic samples are shown in Figure 7 and Figure 8. There was significant pore reduction in the samples fired from 1350°C to 1400°C in agreement with their relative bulk densities (Figure 5). Although the samples fired at 1400°C had the highest density, they also had the largest shrinkage and warping. Pores and ceramic material can be seen in Figure 7a and Figure 7b while micro cracks can be seen Figure 7c. As temperature increases, pores close as the diffusion increases in the ceramic.

Pores and ceramic material can be seen in Figure 7a and Figure 7b while micro cracks can be seen Figure 7c. As temperature increases, pores close as the diffusion increases in the ceramic. These samples presented a proliferation of micro cracks (~400 µm in length) in the bulk of the ceramic body, possibly due to a phase transformation of the ceramic materials at the higher temperature of 1400°C. Thus a firing temperature of 1375°C was selected as appropriate for the ceramic parts.Figure 7. Sectioned ceramic microscopy piece.(A)(B)(C)(D)Figure 8. Optical microscopy (50x) of sintered ceramics at different sintering cycles: a) porosity with mounting media and b) ceramic material, c) micro cracksc3.3.4.X-Ray Diffraction (XRD) analysisBoth sets of the collected XRD data are shown in Figure 9. The figure shows the Bragg peaks used to compare the parts to the machine database to determine the printed and sintered raw material.

The figure shows the Bragg peaks used to compare the parts to the machine database to determine the printed and sintered raw material. The final sintering temperature chosen (1375°C) to fire the ceramic parts for the manufacture of MMCs correlated to a 78% match to the cordierite crystallographic structure found in the literature and was also the highest resulting match. The mismatch in matching the crystallographic structure of the AM built parts to that of cordierite may have been caused by the somewhat inhomogeneous spreading of the mixed raw materials required to make cordierite. Based on such 78% match, the true values of relative bulk density may differ slightly from those presented in Section 3.3.1 due to a change in theoretical density.Figure 9. XRD data from different sintering temperatures of the ceramic samples and software values for cordierite4.

XRD data from different sintering temperatures of the ceramic samples and software values for cordierite4. CONCLUSIONS AND FUTURE WORKThis work presented a novel process to manufacture complex ceramic geometries for the production of metal matrix composites. The overall research goal was to manufacture, test, and evaluate printed cordierite to withstand the environment of metal casting. The process included utilizing a ceramic fabricated by Additive Manufacturing from powder precursor materials, printing, sintering, and post sintering characterization.Traditionally, manufacturers were bound to conventional forming techniques which geometrically limited the design. The use of AM technologies enables the user to manufacture complex ceramic parts that can be tailored for specific applications including MMCs.

The use of AM technologies enables the user to manufacture complex ceramic parts that can be tailored for specific applications including MMCs. This work was focused on the use of AM to form the raw materials used for the production of a cordierite ceramic inserts for MMC production.Additive manufacturing was used to print the precursor powders, alumina (Al2O3), silica (SiO2), and magnesia (MgO), for the production of a cordierite (2MgO•2Al2O3•5SiO2) like material. The correct print parameters (binder selection, layer thickness, heater temperature, binder saturation, spread speed, and drying rate) to successfully print the raw materials were determined experimentally. The materials were characterized pre-and-post sintering to evaluate the properties of the synthetized ceramic.

The materials were characterized pre-and-post sintering to evaluate the properties of the synthetized ceramic. It is recommended there is better control of the initial particle size (>20 micron mean diameter) and shape (from angular to spherical) to increase the flowability of the material during spreading, minimalizing agglomerates from forming and being drug across the build tray. Spherical particles will also provide more uniform initial packing density giving a consistent value for saturation settings.In addition to using raw materials to produce cordierite inserts, the designer may also use printed materials with intricate geometries including those from other raw oxide materials (Al2O3, mullite, enstatite, forsterite, sapphirine spinel, periclase, and corundum) to create ceramics for MMCs. However, the ceramic materials must be able to withstand he drastic thermal shock from the heat and the pressure of the metal matrix being casted.5.

However, the ceramic materials must be able to withstand he drastic thermal shock from the heat and the pressure of the metal matrix being casted.5. ACKNOWLEDGEMENTSThe authors gratefully acknowledge Drs. Brian Lattimer and Patrick Summers of the ExtReMe Lab, the Department of Mechanical Engineering, and the Department of Materials Science and Engineering at Virginia Tech for their resources and support with equipment.6. REFERENCES[1] Clyne, T. W., and Withers, P. J., 1993, An Introduction to Metal Matrix Composites,Cambridge University Press.[2]Suresh, S., Mortensen, A., and Needleman, A., 1993, Fundamentals of Metal-Matrix Composites, Butterworth-Heinemann, Boston.[3]Schleg, F. P., 2003, Technology of Metalcasting, American Foundry Society, Schaumburg, IL.[4]Standke, G., Müller, T., Neubrand, A., Weise, J., and Göpfert, M., 2010, “Cost-Efficient Metal-Ceramic Composites-Novel and Characterisation,” Adv. Eng. Mater., 12(3), pp.

brication of Ceramic and Metal Matrix ComJX>sites From Selective Laser Sintered Ceramic PreformsLucy Deckard and T. Dennis Oaar Lanxide Corporation Newark, DelawareAbstractThis paper will discuss the tool~less fabrication of functional advanced comJX>sites by infusion of a ceramic or metal matrixinto Selective Laser Sintered(SLS) porous ceranU~preforms using Lanxide'spatentedmatrix infusionproc;esses. The fabri~ationofJX>rous preformS of particulate cerami~sby SLS atthe University of Texas at Austin is described in a companion paper. The PRlME}(TI4 pressureless metal infi1trationp~ss was used to infiltrate<alUIninum matrices into both SiC and Al~03 particulate SLS prefonns to make metal matrix comJX>sites without the use of tooling. Also, SiC I~03 ceramic matrix comJX>sites were fabricated using the DIMO}(TI4 directed metal oxiJation process to grow an Al 0 preforms.

Also, SiC I~03 ceramic matrix comJX>sites were fabricated using the DIMO}(TI4 directed metal oxiJation process to grow an Al 0 preforms. Measured properties and microstructures of the resulting composites will be presented and compared to similar comJX>sites made using conventionally fabricated preforms. The rapid prototyping of a SiCiAI MMC electronic power package to near~net shape from an SLS preform will also be describeQ.3 matrix into porous SiC particulate SLS2IntroductionLanxide's matrix infusion processes for fabricating ceramic and metal matrix ~omJX>sites are near~net shape processes in which a matrix is infused into a porous ceramic preform. The final comJX>site shape is dictated by the shape of the preform; therefore, a critical step in the process is the fabrication of a porous preform to near-net shape.

The final comJX>site shape is dictated by the shape of the preform; therefore, a critical step in the process is the fabrication of a porous preform to near-net shape. Prefonns typically have been made using a variety of standard ceramic processing methods, including tape casting, injection molding, green machining, dry pressing, etc. Many of these conventional processes either do not have the complex shape-making capabilities needed or require tooling that is expensive and time-consuming to fabricate for prototypes or small production runs.

The newly-emerging Solid Freeform Fabrication technologies, and in particular Selective Laser Sintering (SLS), appear to be ideally suited to fill the need for a method to fabricate JX>rous preforms to near-net shape quickly and without the use of tooling.One of the objectives of the current work is to evaluate the feasibility of fabricating Ceramic Matrix ComJX>sites (CMCs) and Metal Matrix Composites (MMCs) using preforms formed to net shape using SFF processes. Fabrication of the SLS ceramic preforms is described in a companion paper; the results of work to convert these preforms to CMCs and MMCs via Lanxide's matrix infusion processes are discussed below.Lanxide's Matrix Infusion ProcessesThe DIMO}(TI4 directed metal oxidation process for fabricating ceramic matrix composites (Fig. 1) and the PRIMEXTM pressureless metal infiltration process for metal matrix composites (Fig. 2) In both processes, a ceramic preform is made from the desired reinforcing material, are similar.

1) and the PRIMEXTM pressureless metal infiltration process for metal matrix composites (Fig. 2) In both processes, a ceramic preform is made from the desired reinforcing material, are similar. which can be ceramic particulates or fibers. In the current work, SiC and Al20 3 particles ranging in size from 1 /lm to 60 /lm, which are typical reinforcing materials, were used (fibers were beyond the scope of this work). The preform is formed to the desired shape of the fmal part, a barrier is applied to the upper surfaces to stop infiltration and retain shape, and the preform is placed in contact with the growth alloy which is typically an aluminum alloy. In the CMC process,215the alloy is heated above its melting point to temperatures of 900°C - l0000C in air, and the alloy wicks into the preform while simultaneously oxidizing to form an AL,03 matrix.

This wicking and oxidizing process (referred to as "matrix growth") continues, filling the entire preform until the growth barrier on the top surface is reached The resulting Ceramic Matrix Composite part, consisting of a reinforcement phase and an AI20 matrix with small interconnecting channels of 3 AI, is removed from the alloy. The process for fabricating MMCs is similar, except that the alloy is heated aboveits melting point to temperatures of only 750°C to 850°C in a nitrogen atmosphere; no oxidation ofthe aluminum occurs as it wicks into the preform to form an aluminum matrix composite. No pressure is required to aid infiltration of the alloy into the preform.The chief advantages of these processes are:near-net shape capability (less than 1% shrinkage compared to preform dimensions) • uses comparatively inexpensive raw materials and equipment • no part-specific tooling required after fabrication of preformsExperimental WorkAs described in the companion paper, Marcus, et.

al. supplied Lanxide with both A~03 andSiC particulate SLS preforms made using a variety of sintering parameters. Initially, preforms were supplied in the form of 2" x 2" x 1/4" coupons, then after composition and sintering parameters were established, complex-shaped preforms were fabricated. These preforms were evaluated for compatibility with Lanxide's matrix infusion processes. Initially, preforms were made using either 20 J.1m AI powder or spray-dried PMMA as a binder. We were able to successfully make MMCs and CMCs from both types of preforms; however, the microstructures of the composites made with preforms using the PMMA binder were much more uniform. It was therefore decided that further work would employ the PMMA binder only.Table 1 gives the properties measured to date for MMCs and CMCs made from SLS preforms.

It was therefore decided that further work would employ the PMMA binder only.Table 1 gives the properties measured to date for MMCs and CMCs made from SLS preforms. Although SLS preforms were fabricated from both Al20 and SiC particulate, work concentrated on SiC because both the CMC and MMC demonstration components were to be fabricated using SiC parrlculatereinforcement The reinforcementparticle loadings of the SiC preforms ranged from 4Oto>46 vol%.At least one preform fromeachset(~~imensfabricated using the same SLSParameters) wasiinfiltratedusing Al alloy to make an MMC, and another preform fromthe set was grown to make aCMC.

The SiC preforms generallY infiltrated well; no significant differences in matrix infusion behaviorfor either the CMC orMMCprocesses were observed for specimens made with the various SLSparameters tested The main effect observed in preforms fabricated using differentSLS parameters wa.s variation in green strength which affected ability to withstand handling. Figure 3 shows a microstructure for the MMC; it can>be seen that there is almost no visual evidence oflayering ofthe SiC. However, a relatively large amountof fine porosity (approx. 8vol% as measured by.QIA) was evident, >resulting in.lower thermal conductivityithan would nonnally be expected.

However, a relatively large amountof fine porosity (approx. 8vol% as measured by.QIA) was evident, >resulting in.lower thermal conductivityithan would nonnally be expected. Work is currently concentrating on adjusting infiltration parameters to reduce the occurrence ofthis porosity.3Figure 4 shows a plot of modulus as a function of particleloading for SiC/AI MMCs made at Lanxide using a variety ofpreformingtechniques; it can>beseenth~t:MM:CsfabricatedfromSLS preforms have.elasticirtlodulicomparable>to MMCsmadeusingotlierpref0rmingtechniques, takin~intoaccll~ntthe vol%parti.cle loading.. Figure Sshowsasirnilargraph ofCTEasa function ofloa.ding, and againit·can be seen thatMMCs ma.de using SLS pn;:fonnsbehave.similarly to composites made using conventional preforms.These results indicate that it should be possible to make fully functional MMC components from SLS preforms.

However, many MMC applications require composites with higher reinforcement216and small production runs could be made using SLS preforms, and then if a larger production run is required high volume preforming methods could be used to make essentially the same material.AcknowledgementsThis work was performed under ARPA-sponsored Precompetitive Technology Grant MDA972-92-1-1026. The authors wish to thank Dr. William Coblenz for his encouragement and guidance.GROWTH BARRIEROXIDIZING ATMOSPHERE/PREFORM FORMATIONMATRIX GROWTHREINFORCED CERAMIC COMPONENTFigure 1. A Schematic of the DIMOXTM directed metal oxidation process for making CMCsINFILTRATION ATMOSPHEREINJECTION MOLD PREFORMINFILTRATED MMC PACKAGEFigure 2. A Schematic of the PRIMEXTM pressureless metal infiltration process for making MMCs218Table 1.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFABRICATION OF CERAMIC PARTS USING A DIGITAL LIGHT PROJECTION SYSTEM AND TAPE CASTINGX. Wang, D. De CaussinIndustrial and Manufacturing Engineering Department, California Polytechnic State University, San Luis Obispo, 1 Grand Ave, San Luis Obispo, CA 93407AbstractIn this paper, fabrication of ceramic parts using projection-based stereolithography by curing the mixture of ceramic and photopolymers was investigated. A stereolithography device with a UV LED projection light source was built. A series of resin and ceramic powders, including alumina and zirconia, were experimented to explore the viscosity of mixture and the resultant part quality. It was found that commercial photopolymers are not suitable for this purpose due to the small cure depth.

It was found that commercial photopolymers are not suitable for this purpose due to the small cure depth. A customized photopolymer without photoinhibitor and optimized photoinitator has demonstrated success in producing 3D ceramic green bodies. It was found that the viscosity was very high for all of the mixtures with high ceramic loading which will influence the recoating process. Therefore, tape casting was used to apply the slurry during the build process.1. IntroductionCeramic manufacturing produces parts with desirable traits like high hardness, high temperature strength, and low conductivity. Conventional powder metallurgy approach to manufacture ceramic parts involves high tooling cost and long lead time, which is not desirable for small batch prototyping. Additive manufacturing (AM) provides a fast alternative method to fabricate ceramic parts by building layer by layer from a CAD file.

Additive manufacturing (AM) provides a fast alternative method to fabricate ceramic parts by building layer by layer from a CAD file. Among AM methods, only a few are suitable for producing ceramics, including stereolithography (SL), filament extrusion [1, 2], powder bed fusion, ink writing, and binder jetting [3, 4]. SL process uses ultra-violet radiation from a light source to expose liquid photopolymers to cause cross linking. Under the category of SL, there are three subcategories that describe how the liquid monomer is cured: vector scan, two photon, and mask projection [5]. Digital micromirror device (DMD) based system, invented by SRI International, provides a convenient way to build 3D objects by projecting UV light onto a photopolymer vat using the sliced digital files manipulated by the DMD [6]. A comprehensive of review of ceramic SL can be found in reference [7, 8].

A comprehensive of review of ceramic SL can be found in reference [7, 8]. Ceramic stereolithography with DMD has been investigated using LED as light source [9].In ceramic SL, the viscosity of photopolymer is very important. Lower viscosity is beneficial for the slurry to flow and recoat; however higher viscosity is desired to ensure when the ceramic particles are added, the slurry is stable and there is no sedimentation occurred. The typical particle loading in the slurry is between 40% to 55% volume concentrations [9]. There are two types of suspensions: aqueous and resin. Aqueous suspensions are a mixture of acrylamide and N- N’-methylenebisacrylamide dissolved in deionized water [10]. Since it has a water base, the viscosity is low and flows easily, which makes recoating easy. Even though it has a favorable viscosity, it has many other undesirable traits. An example is that the strength after curing is low which makes it delicate.

Even though it has a favorable viscosity, it has many other undesirable traits. An example is that the strength after curing is low which makes it delicate. It also doesn’t hold the ceramic particles in suspension well because of the low viscosity. The aqueous suspension needs to constantly be agitated to keep the particles suspended. Research from Griffith and Halloran have shown that aqueous acrylamide suspensions have very good fluidity but the depth of cure is not optimal due to the refraction from the powders914[9]. In contrast to the aqueous suspension, resin based suspensions create a much stronger green body after curing but also have a higher viscosity [10]. This type is the common commercial resin used in stereolithography. Brady and Halloran have researched that 200 mPa·s is the lowest viscosity that will ensure the ceramic particles will be stable against sedimentation [11]. Cure depth is another important factor for ceramic SL.

Cure depth is another important factor for ceramic SL. The cure depth of the mixture of resin and ceramic powders is dependent on the species of the photopolymer, the ceramic loading, and the UV exposure dose. The ceramic particles change the way the light penetrates and can either scatter or absorb to change the refractive index. Modified Beer-Lambert law has been proposed to model the cure depth in ceramic SL [9]. wayThis paper presents a DMD based ceramic SL system that exposes the slurry made of ceramic powders and photopolymers to build green body followed by debinding. A DMD projection system was built in-house for this study. Both commercial resin and customized resin were investigated. The viscosity of the slurry with were characterized.2. EXPERIMENT2.1 EQUIPMENT SETUPThe exposure system consists of a 5.5 W LED UV light source, a motherboard with DMD from Texas Instruments, and optics (Wintech Lightcrafter PRO4500, Wintech Inc. TX).

TX). The UV LED releases 405nm wavelength of light in a field of view of 65.6mm × 41mm. The DMD has a resolution of 912 × 1140 (WXGA) with diamond pixel. The focal distance is 92mm. The control board communicates with a computer through the Texas Instruments provided graphical user interface (GUI) that controls the image upload and exposure time with a HDMI interface. The STL file was sliced using Freesteel Slicer and each individual slice was uploaded to the firmware and stored on the external boards. Exposure time of each layer is controlled by a computer connected to the control board.The platform system moves the build downwards in the z direction between each exposure. A XY-stage was used to control the movement. The resolution in the z direction was 10 microns. The platform is made of an L shaped sheet aluminum that the resin adheres to and was attached to the XY- stage.

The resolution in the z direction was 10 microns. The platform is made of an L shaped sheet aluminum that the resin adheres to and was attached to the XY- stage. The platform was merged in a vat filled with photopolymer.The recoating was achieved using an eccentric motor to provide vibration which would cause the resin to flow off to cover the finished layer. When the ceramic material was applied to the resin, it made the viscosity increase significantly. It was found that it is difficult for the slurry to flow over. Instead, a tape casting method with squeegee and a template were used for recoating. The final configuration is shown in Figure 1.915Figure 1 Final stereolithography 3D printer with all subcomponents mounted2.2 EXPERIMENTS2.2.1 MATERIALSThe exposure was done with the instrument developed in section 2.1. Two resins have been used in the experiments: a commercial photopolymer (AlphaSense Inc. DE), and a customized resin (Spectra Group Ltd., Millbury OH).

Two resins have been used in the experiments: a commercial photopolymer (AlphaSense Inc. DE), and a customized resin (Spectra Group Ltd., Millbury OH). Three powders were investigated: alumina with particle size of 3.0 µm (Hudson Supply, OH), zirconia with particle size of 40 nm (TZ-3Y-E, Tosoh Corp. Japan), and alumina with particle size of 500 nm (Ceralox APA-0.5, Sasol Gmbh).2.2.2 PROCESSThe powders were mixed with resin in a container first and then transferred to the vat which is underneath the projection lens. The viscosity of the slurry was measured using a viscometer (Brookfield DV-I). The sliced layer image was sent to the projector and exposed the slurry. After one layer was exposed, the platform lowered down and a blade was used to recoat another layer until all layers were completed.

After one layer was exposed, the platform lowered down and a blade was used to recoat another layer until all layers were completed. The green body was removed from the platform and then debinded in a furnace (SunFire 10, Neytech Inc).A series of experiments have been conducted: (1) commercial resin only to test the functionality of the instrument; (2) commercial resin with 3.0 µm alumina and 40 nm zirconia; (3) customized resin with 500 nm alumina powders. The experimental details are described in section 2.2.3 to 2.3.5.2.2.3 COMMERCIAL RESIN ONLY EXPERIMENTThe first experiment was done to validate the system functionality with resin only. A commercial transparent yellow photopolymer (AlphaSense Inc. DE) was used. The viscosity of the resin was measured using a Brookfield DV-I viscometer, with a cylindrical spindle. The viscosity was found to be 37.6 cP at room temperature.

DE) was used. The viscosity of the resin was measured using a Brookfield DV-I viscometer, with a cylindrical spindle. The viscosity was found to be 37.6 cP at room temperature. The exposure was 20 seconds and the recoating time was 40 seconds where the plate would step down and the resin would flow to recoat the plate. Simple rectangle and oval shapes were used as the test parts with a 0.1mm slice thickness. During the exposure, the liquid resin visibly cured when the image was projected.2.2.4 COMMERCIAL RESIN WITH ALUMINA AND ZIRCONIA916The first test with alumina and the AlphaSense resin slurry has a particle loading of 50% by volume. The alumina powder has a particle size of 3.0 micron (Hudson Supply, OH). This mixture resulted in a very powdery end product that was not holding together which is not suitable for the exposure. The second batch has a 40% particle loading by volume.

This mixture resulted in a very powdery end product that was not holding together which is not suitable for the exposure. The second batch has a 40% particle loading by volume. The viscosity of the slurry was measured to be around 100,000 cP.For the zirconia mixture, a lower volume percentage of 26% particle loading by volume was used to ensure the right viscosity suitable for recoating. The Zirconia has a particle size of 40 nm (TZ-3Y-E, Tosoh Corp. Japan) which is expected to yield a deeper cure depth.2.2.5 CUSTOMIZED RESIN WITH FINER POWDERSThe cure depth from the experiment in section 2.2.2 is not enough with current settings. To increase the cure depth the resin and ceramic material were changed. A custom resin was made for this purpose by Spectra Group Ltd. (Millbury OH). The resin is an acrylic monomer base with a viscosity of 85 cP at 25ºC. It has no photo inhibitors added in order to maximize the cure depth.

(Millbury OH). The resin is an acrylic monomer base with a viscosity of 85 cP at 25ºC. It has no photo inhibitors added in order to maximize the cure depth. When this resin is exposed to UV light by itself, it can reach a cure depth of up to one inch depending on the light source power.The alumina powder has a particle size of 500 nm (Ceralox APA-0.5, Sasol Gmbh). The alumina has 500 ppm magnesia as dopant to help minimize agglomeration and maximize dispersion. The small particle size and magnesia were chosen to help keep the viscosity lower once it is added to the resin. Two batches of slurry were prepared: 35% and 40% particle loading.2.3 DEBINDINGThe green parts were sintered in a furnace (SunFire 10, Neytech Inc.). This has different ramp features and holds that are programmable in order to get the desired sintering process. The first phase for our test was the resin burnout. For this, the furnace was ramped up at 25°C/min to 732°C and held for 20 minutes.3.

The first phase for our test was the resin burnout. For this, the furnace was ramped up at 25°C/min to 732°C and held for 20 minutes.3. RESULTS AND DISCUSSION3.1 VISCOSITY OF SLURRYThe viscosity of the slurry is important and will influence the recoating and stability of the slurry. As discussed earlier, in order to have a successful SL process with vat recoating, the viscosity needs to be below 3000cP to flow properly and above 200cP to limit sedimentation. A rheometer (DHR-2, TA Instruments) was used to characterize the viscosity. All testing was done at 25ºC. The flow sweep tests the viscosity at different shear rates taking the average of the measurements over 60 seconds for each data point. The range of shear rates varied between 0.01 sec-1 and 10 sec-1. This lower shear rate range was used because it is close to the conditions during SL.

The range of shear rates varied between 0.01 sec-1 and 10 sec-1. This lower shear rate range was used because it is close to the conditions during SL. The results are show in Figure 2.The pure customized resin (from Spectra Group) exhibits Newtonian fluid behavior because the viscosity remains relatively constant with a changing shear rate. It shows a viscosity around 100cP which is close to the manufacturer’s specification of 85cP. All of the samples with 500 nm alumina powders (from Ceralox) however seem to show a shear thinning behavior which means that as the shear rate is increased, the viscosity decreases. This is probably due to the alumina being a suspension in the resin rather than a complete uniform dispersion. One important point in Figure 2 is where the viscosity passes the 3000cP measurement. This is between the 10%917and 20% particle loading which is much lower than the optimal loading.

One important point in Figure 2 is where the viscosity passes the 3000cP measurement. This is between the 10%917and 20% particle loading which is much lower than the optimal loading. A loading closer to 50% is desirable in order to produce a green body that can be sintered to full densification. This means that for the Ceralox alumina and resin mixture, a vat recoating system will not work because the viscosity gets very high with the higher loading percentages. In order to resolve this, the recoating will use a tape casting method so the alumina particles can be loaded to a higher percentage. Another observation when mixing these different test samples was the amount of sedimentation at different levels. After mixing the compounds and letting them settle for a minute, there was a very thin layer of alumina at the bottom. Once the mixture got to about 30% particle loading there was no noticeable sedimentation.

Once the mixture got to about 30% particle loading there was no noticeable sedimentation. This is due to a flocculated powder bed that entered the compaction phase of powder sedimentation. When the particle loading went above 40%, the viscosity was too high to be measured. It was a very thick sticky peanut butter consistency and would hold its shape without flowing.Figure 2 Viscosity test of customized resin with alumina in varying volume percentages3.2 CURE DEPTHThe curing of pure resin (AlphaSense resin) has been cured by 20 seconds exposure. During the exposure, the liquid resin visibly cured when the image was projected. Even though the UV light was on for 20 seconds, the layer was cured in about 5 to 10 seconds.

During the exposure, the liquid resin visibly cured when the image was projected. Even though the UV light was on for 20 seconds, the layer was cured in about 5 to 10 seconds. The extra exposure time does not affect the final part; it just slows down the print time.However with the slurry made of 40% 3.0 micron alumina, after the first exposure of 40 seconds (trying to get the maximum cure depth), the result was a thin film of about 50 microns. This cure depth was not sufficient because it did not allow subsequent layers to bond to the previous one. This resulted in the parts being pushed around as each layer was applied which ultimately would not yield a viable part. When exposure time was increased, the cure depth did918not increase. This was because the reflection of particles reduced the UV light absorption.

When exposure time was increased, the cure depth did918not increase. This was because the reflection of particles reduced the UV light absorption. The grey color of the alumina powder itself also partially contributes to the less absorption.The cure depth of this zirconia mixture was much higher than the alumina at about 0.25mm and each subsequent layer was successfully bonding with the previous. From exposure point of view this is desired but from sintering perspective, the particle loading was too low at 26%, which would not be sufficient to provide a fully ceramic part after sintering.In order to successfully obtain a green body with reasonable relative density and also ensure the cure depth is enough, the resin must be optimized. The commercial resin usually has photoinhibitor, which is to prevent excessive curing depth in conventional SL. For the purpose of curing mixture of particle and resin, the concentration of photoinhibitor must be reduced.

For the purpose of curing mixture of particle and resin, the concentration of photoinhibitor must be reduced. In the customized resin, the photoinhibitor was completed removed to maximize the cure depth. At the same time, the photoinitator must also be optimized to increase photo polymerization. The percentage of photoinitiator (H-NU 400 IL, Spectra Group, OH) was optimized for the 405 nm wavelength. Figure 3 shows the absorption with different concentration. The absorption of the resin used in this study was about 7%. Smaller particle size also helps to increase cure depth; therefore, alumina powders with finer particle size (500 nm, Ceralox) was chosen to make slurry with this customized resin.Figure 3 Absorption of resin with different concentration of photoinitiator [courtesy of Spectra Group]In order to test the cure depth of this mixture, a particle loading of 35% was printed.

The step depth between each exposure was 0.2mm and a tape casting method was used to apply the slurry. Rather than dipping it into the bath, the resin mixture was squeegeed onto the plate. Each exposed layer adhered to the previous layer so the cure depth was sufficient with this resin. The resultant green body has low relative density that handling is difficult. Therefore another batch with particle loading increased to 40% was prepared.919Figure 4 40% particle loading of alumina with customized resinFigure 4 shows the final green part with 40% concentration of Ceralox alumina. Each layer has very high accuracy with regards to the features. The small disconnect between the “C” and “P” was kept separate throughout the build. The one discrepancy visible is the change between each height. As the platform lowers, the layer height gets thicker in some places and sometimes didn’t settled directly on the previous layer.

The one discrepancy visible is the change between each height. As the platform lowers, the layer height gets thicker in some places and sometimes didn’t settled directly on the previous layer. The thicker portions are due to the tape casting and not getting a perfectly smooth layer application. The disconnection between layers is due to the platform moving between layers while building. In order to resolve this, another bolt can be added to prevent any rotation on the printing plate. Another experiment was conducted with this improvement. Three different parts with different sizes were successfully exposed to examine the feature’s fidelity with part size. The results is shown in Figure 5. Each subsequent layer had very good placement on the previous which confirms that the print plate no longer has rotational movement to take away from the accuracy. The parts are a quarter inch tall and the detail was held through each layer. The smallest mustangs produced fine details.

The parts are a quarter inch tall and the detail was held through each layer. The smallest mustangs produced fine details. The gap between the front hoofs is 0.01 inch and it was held separate in all the builds which shows very good layer accuracy.Figure 5 Exposure with different feature sizeFigure 6 shows the part during resin burnout at approximately 400°C. The black is the resin carbonizing and burning out of the part. After it reached 732°C, the carbonizing was over and the part went back to white showing that the resin was gone.Figure 6 Resin burnout initial stages showing carbonization4. CONCLUSIONS9201. Higher viscosity is desired to prevent particle sedimentation during ceramic SL, but make it difficult to recoat. Tape casting method demonstrated success with good layer uniformity. 40% particle loading by volume has been achieved in this study which yields higher green body density for post sintering.2.

Tape casting method demonstrated success with good layer uniformity. 40% particle loading by volume has been achieved in this study which yields higher green body density for post sintering.2. Commercial photopolymers are not suitable for ceramic SL due to the cure depth is too low. The resultant green body doesn’t have enough adhesion between layers.3. A customized photopolymer without photoinhibitor and optimized photoinitiator, has been demonstrated successfully producing alumina ceramic green body with very good resolution. ACKNOWLEDGEMENTSThe support from California Polytechnic State University at San Luis Obispo’s RSCAfunding is greatly appreciated.REFERENCES1 Bandyopadhyay, A., Panda, R., Janas, V., Agarwala, M., Danforth, S., & Safari, A. (1997). Processing of Piezocomposites by Fused Deposition Technique. Journal of the American Ceramic Society, 80(6), 1366-1372. 2 Bandyopadhyay, A., Panda, Mcnulty, Mohammadi, Danforth, & Safari. (1998).

brication of Complex 3D Micro-Scale Scaffolds and Drug Delivery Devices using Dynamic Mask Projection Microstereolithography1Jae-Won Choi, 2In-Baek Park, 1Ryan Wicker, 2Seok-Hee Lee, 3Ho-Chan Kim 1W.M.Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, TX 79968 2School of Mechanical Engineering, Pusan National University, San 30 Jangjeon-dong, Geumjeong-gu, Busan 609-735, S. Korea 3School of Mechanical Engineering, Andong National University, 388 Songchun-dong, Andong-si, Gyeongbuk, 760-749, S. KoreaReviewed, accepted September 10, 2008AbstractMicrostereolithography (μSL) technology can fabricate three-dimensional (3D) tissue engineered scaffolds with controlled biochemical and mechanical micro-architectures. AμSL system for tissue engineering was developed using a Digital Micromirror Device (DMDTM) for dynamic pattern generation and an ultraviolet (UV) lamp filtered at 365 nm for crosslinking the photoreactive polymer solution.

The μSL system was designed with x-y resolution of ~2 μm and a vertical (z) resolution of ~1 μm. To demonstrate the use of μSL in tissue engineering, poly(propylene fumarate) (PPF) was synthesized with a molecular weight of ~1200 Da. The viscosity of the PPF was reduced to ~150 cP (at 50 oC) by mixing with diethyl fumarate (DEF) in the ratio of 7:3 (w/w). Finally, ~2 % (w/w) of (bis(2,4,6- trimethylbenzoyl) phenylphosphine oxide (BAPO) was added to the solution to serve as a the curing photoinitiator. characteristics of the synthesized PPF, and the resulting system and photopolymer were usedCure depth experiments were performedto determineto construct a variety of 3D porous scaffolds with interconnected pores between 100 and 150μm and a micro-needle array with height of ~800 μm and individual tip diameters of ~20 μm.

SEM and microscope images of the micro-architectures illustrate that the developed μSL system is a promising technology for producing biodegradable and biocompatiblemicrostructures.INTRODUCTIONIn 1993, microstereolithography (μSL) technology, which evolved from conventional stereolithography, was suggested by Ikuta and Kirowatari [1] and Takagi and Nakajima [2]for producing micro-scale complex structures. The first type of μSL machine was based on the vector-based scanning SL method, referred to as line-scan. In 1996, Nakamoto andYamaguchi [3] suggested a mask-based μSL system, where entire layers were projected at a single time using a physical mask instead of scanning the surface with a laser beam. In 1997,an LCD (Liquid Crystal Display)-based projection μSL system was developed by Bertsch et al. [4]. In 2001, Bertsch et al. [5] enhanced the resolution limitation of the LCD system byreplacing the LCD with a DMDTM (Digital Micromirror Device).

[4]. In 2001, Bertsch et al. [5] enhanced the resolution limitation of the LCD system byreplacing the LCD with a DMDTM (Digital Micromirror Device). In μSL, the LCDs and DMDs are used as dynamic pattern generators or dynamic masks that create any desired pattern on the resin surface, and thus, these methods are referred to as dynamic maskprojection methods as opposed to the scanning method described previously. Since the firstμSL system was suggested, many researchers have developed similar μSL systems because of the superior capability for producing 3D complex microstructures [6–14].These μSL systems can produce these scaffoldsimplantable scaffolds using photocurable to be biodegradable and biocompatible.biomaterials Biodegradability means that the scaffold has to be chemically and gradually degraded in vivo, the scaffold has to be destructed without causing cytotoxicity leaving the desired shape of the regenerated organ or tissue.

Biocompatibility also means that the surface of the scaffold hasthat allowto be chemically compatible such that cells attach to and grown on the scaffold, and it is imperative that the scaffold does not induce any undesired reactions (or immune responses) with neighboring organs or tissues [15]. Fortunately, biodegradable and biocompatiblematerials can be synthesized for use in μSL. Cooke et al. [16] first demonstrated a poly(propylene fumarate) (PPF) scaffold that was fabricated in a conventional SL system, but the scaffold micro-architecture was limited because of the resolution of the system. Lee et al. [17, 18] fabricated porous PPF scaffolds using scanning microstereolithography (µSL) and conducted a cell study using the fabricated scaffolds. The scaffolds fabricated by Lee et al.,although fabricated using µSL, were of simple shapes and provided no real 3D architectures.

The scaffolds fabricated by Lee et al.,although fabricated using µSL, were of simple shapes and provided no real 3D architectures. As a result, the capability of µSL to fabricate complex 3D micro-architectures has yet to be fully explored, including examining and optimizing the effects of micro-architectures on cell behavior.In addition to scaffold fabrication, µSL systems can produce biocompatible drug delivery devices using the biomaterials. Micro-needles, for example, are drug delivery devices that can be designed and manufactured using µSL to puncture the skin without pain. The punctured micro-holes in the skin can be utilized as a path for drugs, which has evenhigh molecular weight such as a protein. Many researchers have produced micro-needles using microelectromechanical system (MEMS) and lithography, electro-forming, and molding (LIGA) technologies [19–24], although these micro-needles were also limited on the geometric complexity by the manufacturing technology.

By using µSL, it may now bepossible to optimize micro-needle design in terms of puncture force, drug delivery rate, and other factors.This paper demonstrates development of a DMDTM (Digital Micromirror Device)- based microstereolithography (μSL) system, PPF synthesis and characterization, and fabrication of 3D micro-scaffolds and micro-needles using commercial photopolymers and PPF. PPF was synthesized using DEF (Diethyl fumarate) and PG (Propylene glycol) alongwith the hydroquinone and zinc chloride. The DMD-based μSL was developed with thelateral resolution of ~2 μm, and vertical resolution of ~1 μm. The synthesized PPF was mixed with the DEF to reduce the PPF viscosity for use in the µSL system.

The synthesized PPF was mixed with the DEF to reduce the PPF viscosity for use in the µSL system. In order to characterize PPF/DEF, curing, penetration depth and critical energy studies were performed.Finally, the fabricated microstructures were examined by SEM and optical microscopy toexamine the ability of the μSL system to produce complex micro-scaffold and micro-needle geometries. The following sections describe these experiments and their results in more detail.DYNAMIC MASK PROJECTION MICROSTEREOLITHOGRAPHYComponents and Principle In microstereolithography (µSL), the fabrication method works in the same fashion as in SL.That is, 3D microstructures can be produced by slicing a 3D model with a computer program, solidifying, and stacking images layer by layer in the system [8]. The achievable resolution in µSL is within 10 µm, whereas it is several tens microns in conventional SL.

The achievable resolution in µSL is within 10 µm, whereas it is several tens microns in conventional SL. There are mainly two types of µSL; one is scanning-based µSL using a focused laser spot and the otheris projection-based µSL using a projected light pattern. In scanning µSL, the laser spot is focused and scanned onto the resin surface using a XY-stage to move the spot instead ofgalvanometer mirror, which induces a defocusing problem [1, 25]. In projection μSL, an illuminated light is patterned using a high-resolution pattern generator, and projected onto theresin surface. Projecting the entire image onto the surface results in solidifying the entire layer at a time without having to scan the surface [4, 6, 7].

Projecting the entire image onto the surface results in solidifying the entire layer at a time without having to scan the surface [4, 6, 7]. In this work, DMD-based dynamic mask projection µSL was used, which was developed previously [8, 9, 12, 14].DMD-based dynamic mask projection µSL consists of the light emission subsystem(a lamp, an optical fiber, a filter, and a collimating lens set), the light delivery subsystem (a LightGateTM, a tube lens, and a reflecting mirror), the dynamic pattern generation subsystem (the DMDTM), the image focusing subsystem (a modular focusing unit and an objective lens), and the build subsystem (a Z-stage, a platform, a resin vat, and a hot plate) as shown in Fig. 1[8, 14].Fig. 1 Developed dynamic mask projection μSL systemIn the light emission subsystem, the 200 W mercury lamp (Omnicure S2000TM, EXFO, Co., Canada) containing the filter with the wavelength of 365 nm was selected for the light source.

The emitted light was delivered via the optical fiber. The collimating lens set,which consists of two convex lenses, was used for collimating the light and illuminating to the DMDTM.In order to make the light delivery subsystem compact, the light path must be reduced. To do this, the LightGateTM (Unaxis Co., USA), which was specially coated for UV light was used as a prism. The LightGateTM reflects the light to DMDTM and transmits the returned pattern to the tube lens. To deliver the patterned light toward the objective lens, the tube lens (Achromat doublet lens, MellesGriot Co., USA) with the focal length of 120 mm and reflecting mirror were used. The tube lens played a role in collimating the patterned light, and was positioned at the focal length from the DMDTM.

The tube lens played a role in collimating the patterned light, and was positioned at the focal length from the DMDTM. The reflected light from DMDTM was assumed as the source, therefore the light between the tube lens and the objective lens could be collimated.In the pattern generation subsystem, the DMDTM (DMD Starter Kit, Texas Instruments, USA) was chosen for a dynamic pattern generator. It consists of ~786,000 micromirrors (1024 × 768), in which each mirror is 13.68 µm along each side, and is independently tilted at ±12 degrees by an electrostatic force. The principle of generating the light pattern using DMDTM is that the incident light is reflected in two directions according to the mirror tilt angles, and one of the reflected light bundles is the pattern. The bundle of the light reflected at the +12 degrees makes the desired pattern, which is projected on the resin surface through the tube lens and objective lens, whereas the other bundle is projected to a dummy direction.

The DMDTM makes a certain pattern by tilting each mirror at ±12 degrees according to the binary information of each pixel (one pixel is the same as one micromirror), once the binary image is transferred to DMDTM board.The binary image is generated from the sliced section, which consists of point data with at least one loop. An example of the process for creating the binary image is providedin Fig. 2. In the case where there is more than two loops, there is not any topology relationship between the loops. Therefore, the topology in each section has to be found, and the binary image can be generated according to the number of the loops surrounding a certain loop. That is, each loop is painted black if the number of surrounding loops including itselfis odd. Otherwise it is painted white. Each section is painted from the outer loop to the inner loop. In Fig. 2, the loops L1, L2, L3 and L4 have the number of the surrounding loops (including itself) of 1, 2, 2, and 3, respectively.

Each section is painted from the outer loop to the inner loop. In Fig. 2, the loops L1, L2, L3 and L4 have the number of the surrounding loops (including itself) of 1, 2, 2, and 3, respectively. Thus, L1 is first painted black, then L2 and L3 are painted white, and then the L4 is painted black.In the image focusing subsystem, two kinds of objective lenses with numerical aperture (N.A.) of 0.13 and 0.3 (CFI Plan Flour, Nikon, Japan) were selectively used as a projection lens. The focal lengths (fo) of the objective lenses with the N.A. of 0.13 and 0.3are 50 mm and 20 mm, respectively. The objective lenses were selectively used according to the reduction ratio, which determines the achievable resolution and fabrication volume.Fig.

The objective lenses were selectively used according to the reduction ratio, which determines the achievable resolution and fabrication volume.Fig. 2 Binary image generation: (a) loops in a layer, (b) L1 is painted black (L1 is the outer most loop), (c) L2 and L3 are painted white, and (d) L4 is painted black (the final binary image) In addition, the modular focusing unit with the resolution of 1 µm was used for adjusting the distance between the resin surface and objective lens. To simulate and optimize the designed optical system, ZEMAX software was used. The distances between the DMDTM surface and the tube lens (ldt), and the tube lens and the objective lens (lto) as shown in Fig. 3, were optimized. Table 1 shows the original and optimized distances. The magnification can be calculated by the ratio of the distances ldt and fo by the basic lens equation, because the light between the tube lens and the objective lens was collimated.

The magnification can be calculated by the ratio of the distances ldt and fo by the basic lens equation, because the light between the tube lens and the objective lens was collimated. Therefore, the magnifications using two objective lenses were ~0.434, and ~0.174, so that 1 pixel (13.68 µm) on the DMDTM would represent ~5.9 µm and ~2.4 µm, respectively.Fig. 3 The distances among DMDTM, tube lens, objective lens, and resin surfaceTable 1 The original and optimized distances between the opticsObjective Lensldt (mm)Original distanceOptimized distancelto (mm)Original distanceOptimized distanceMagnification (fo/ ldt)N.A. 0.13120115.2150203.350/115.2 (0.434)N.A. 0.3120115.2150186.520/115.2 (0.174)The build subsystem consisted of a Z-stage with the resolution of 100 nm for stacking, a platform for attaching a substrate, a vat for containing the liquid solution, and a hot plate for controlling the solution temperature.

The Z-stage makes a new solution surfacewith the desired layer thickness by moving downward deeply and then moving upward to the predetermined position. That is, the previously cured layer is immersed deeply so the neighboring solution flows on the top of the layer, and once returned to the predetermined position, the new uncured resin layer is allowed to settle for a certain settling time. Thesettling time depends on the solution viscosity and can be experimentally determined. Fig. 4 shows the process of refreshing the solution surface. Because the recoating process involvesresin flow, the viscosity is a dominant factor in μSL, and a low viscosity solution is needed to ensure precise layer thicknesses and fast fabrication by reducing settling times.Fig.

4 Building process: (a) irradiation and curing, (b) moving downward deeply for the neighboring solution to flow on the top, (c) moving upward to the desired position, and (d) waiting until the solution surface becomes uniformPPF/DEF PREPOLYMERPPF Synthesis Diethyl fumarate (DEF), propylene glycol (PG), zinc chloride, hydroquinone were purchasedfrom Fisher Scientific Korea (Seoul, S. Korea), and bis(2,4,6-trimethylbenzoyl) phenylphosphine oxide (BAPO) was purchased from Ciba Specialty Korea (Seoul, S. Korea). Poly(propylene fumarate) (PPF) was synthesized with a two-step reaction as previously reported [26–30]. Briefly, DEF and PG were prepared as main components with the molarratio of 1 : 3 along with the zinc chloride (0.01 mol) as a catalyst, and hydroquinone (0.002 mol) as a crosslinking inhibitor.

Briefly, DEF and PG were prepared as main components with the molarratio of 1 : 3 along with the zinc chloride (0.01 mol) as a catalyst, and hydroquinone (0.002 mol) as a crosslinking inhibitor. In the first reaction, the prepared solution was mixed by an overhead mechanical stirrer, and heated from 100 oC to 150 oC with an increase of 10 oC every 20 minutes under nitrogen. It was maintained for 7 hours (the time is equivalent to ~90 % of reaction) once the temperature reached 150 oC. Diester intermediates as a main product, and ethanol as a byproduct were produced through transestrification at this time. In the second reaction, the solution was maintained at 130 oC under vacuum for 1 hour (the time is equivalent to the predetermined molecular weight of the PPF). PPF as a main product and PG as a byproduct were produced.

PPF as a main product and PG as a byproduct were produced. Synthesized PPF was not purified because it wasassumed that any purification process may not affect PPF manufacturability in the μSL system.PPF CharacterizationsTo verify the utilization of synthesized PPF in the developed μSL system, the viscosity has to be measured and adjusted. If the solution is too viscous (usually the viscosity of thesolution in μSL is maintained below ~200 cP), the solution surface cannot be easily refreshed because of the building process (see Fig. 4). The viscosity of the synthesized PPF was measured using a viscometer (SV-10, AND Co., Japan) and gel permeation chromatography (GPC) was used to measure its molecular weight. The measured viscosity was ~12000 cP at 47 oC as shown in Fig. 5, and the average molecular weight was determined to be ~1200 Da.Fig.

The measured viscosity was ~12000 cP at 47 oC as shown in Fig. 5, and the average molecular weight was determined to be ~1200 Da.Fig. 5 PPF viscosity according to temperaturePPF/DEF prepolymer preparation The viscosity of the PPF depends on the time of the second reaction, and the viscosity of thesolution has to be modified for μSL as previously mentioned. Therefore, a diluent is neededto lower the viscosity of the solution without changing the desired biomaterial properties. DEF, which is a main material for PPF synthesis, was used as the diluent as it has been reported that DEF does not affect the biomaterial properties of PPF. In addition, DEF has acarbon double bond, so it participates in crosslinking [29]. In this work, DEF was added to PPF with the ratio of 3:7 (w/w), and the ratio was obtained from the work of Fisher et al. [29]. Fisher et al.

In this work, DEF was added to PPF with the ratio of 3:7 (w/w), and the ratio was obtained from the work of Fisher et al. [29]. Fisher et al. showed that the mixture of DEF/PPF (~3:7 by weight%), where the molecular weight of the synthesized PPF was 1260 Da, had the highest elastic modulus and the fracturestrength. The viscosity of the PPF/DEF as a function of temperature is shown in Fig. 6 and was measured using the same viscometer as above. The PPF/DEF prepolymer was prepared by stirring for 12 hours along with the 2% (w/w) of bis(2,4,6-trimethylbenzoyl) phenylphosphine oxide (BAPO) as a photoinitiator. In Fig. 6, the viscosity of the PPF/DEFrepresents ~700 cP at room temperature. This viscosity is relatively high compared to desired value of ~200 cP. However, the viscosity of the solution was, logarithmically reduced by elevating the temperature, and the viscosity was ~150 cP at 50 oC.

However, the viscosity of the solution was, logarithmically reduced by elevating the temperature, and the viscosity was ~150 cP at 50 oC. This value is suitable for μSL [8], so a hot plate was installed under the vat to maintain the temperature of the solution at 50 oC.Fig. 6 PPF/DEF viscosity according to temperatureCure depth experiment A photocurable solution is crosslinked by connecting a monomer, oligomer or polymer chain, where they have carbon double bonds that can be broken down by a radical. The photons from the projected light break down the photoinitiator into a radical along the penetrationdirection of the light. The radicals then bond the neighboring monomer, oligomer, or polymer by breaking carbon double bonds.

The radicals then bond the neighboring monomer, oligomer, or polymer by breaking carbon double bonds. Therefore, in terms of 3D microfabrication, the penetration depth of the light and critical energy at photoinitiation are important and need tobe controlled.To examine the penetration depth and critical energy, curing depth experiments were conducted using the PPF/DEF prepolymer. The energy delivered on the solution surface (Emax) penetrates into the solution. The energy inside the solution at the depth ‘z’ (E(z)) is defined by Beer-Lambert law as described in Eq. 1 below [31], where Dp is the penetration depth of the solution. By introducing the critical energy (Ec) into the Eq. 1, the curing depth (Cd) can be defined as in Eq. 2, where Ec is the energy at the gel point. The gel point is the point at which solidification begins. Therefore, two important characteristics of the photocurable solution are Ec and Dp.

2, where Ec is the energy at the gel point. The gel point is the point at which solidification begins. Therefore, two important characteristics of the photocurable solution are Ec and Dp. These can be experimentally determined through measuring of the curing depth according to the exposure energy. From the determined values of Ec and Dp, the exposure energy and stacking thickness can be chosen. In addition, it is important to note that the smaller the curing depth, the ability to fabricate down-facing and complex microstructures is improved.To conduct the curing depth experiment, the curing model as shown in Fig. 7 was used. The curing model consists of posts, which were stacked with 10 layers (1 mm long)with the layer thicknesses of the 100 µm, and crossbeams, which were fabricated in the lastlayer with the given exposure energy. The irradiance in the developed μSL was 33.8 mW/cm2, and the exposure energy was controlled by opening the shutter for 1, 2, 3, 4, and 5 s.

The irradiance in the developed μSL was 33.8 mW/cm2, and the exposure energy was controlled by opening the shutter for 1, 2, 3, 4, and 5 s. The thicknesses of the 4 crossbeams were measured at the center using a microscope (DFC280, LEICA, Germany).zE )(=Emax⋅e(−z pD)(1)Cd=Ez (max)=Dp⋅ln(E max E c)(2)Fig. 7 3D specimen model for curing depth experimentThe specimens for curing characteristics were fabricated as shown in Fig. 8. The crossbeams for 1 s exposure time (= 33.8 mJ/cm2) were not obtained because the thickness of each individual crossbeam was so small that it was broken during rinsing. Fig. 9 represents the cure depth graph by measuring the thickness of the center of the crossbeams. Accordingto Fig. 11 and Eq. 2, the critical energy and penetration depth were determined to be 32.4 mJ/cm2 and 78 µm, respectively.(a)(b)(c)(d)Fig.

Accordingto Fig. 11 and Eq. 2, the critical energy and penetration depth were determined to be 32.4 mJ/cm2 and 78 µm, respectively.(a)(b)(c)(d)Fig. 8 Fabricated specimens with the crossbeams: (a) 2 s, (b) 3 s, (c) 4 s, and (d) 5 s exposure timeFABRICATED MICROSTRUCTURESFabricated microstructure using commercial photopolymersCommercial photocurable resins are candidates for use in μSL, because many resins are sufficiently reactive with UV light wavelengths used in μSL. Alternatively, commercial monomers along with a photoinitiator can also designed and used for specific purposes.Using the developed μSL system, several microstructures with the complex features were fabricated as shown in Fig. 10. Each microstructure was fabricated using differentfabrication conditions and materials as shown in Table 2. Fig. 10 (a) ~ (d) show SEM images of a micro-wineglass, micro-cup, micro-bishop, and micro-springs, respectively.

10ModelTotal layernumberLayerthickness (µm)ObjectivelensMaterial(a)(b)3002005120.30.13SI40a: IBXAb = 1 : 1 (w/w)(c)(d)1343002040.30.3IBXA: HDDAc: BEDd = 8 : 1 : 1 (by wt%) DMPAe 5 % (w/w) as a photoinitiatoraSI40 was purchased from 3D systems (USA), and suitable for conventional SL systembIBXA, cHDDA, and dBED represent Isobornyl acrylate, 1,6-Hexanediol diacrylate, Bisphenol-A-ethoxylated (4) diacrylate,respectively, and purchased from Woorim Chem Tech Co., S. Korea, Miwon Commercial Co., S. Korea, Hannong ChemicalsCo., S. Korea, respectivelydDMPA represents Dimethoxy-2-phenylacetophenone, and purchased from Fisher Scientific Korea, Seoul, S. KoreaFabricated microstructures using PPF/DEF Scaffolds were also fabricated as shown in Fig. 11 ~ Fig. 13. Fig. 11 shows a kidney scaffold with a volume of ~1400 µm × 820 µm × 700 µm and pore size of ~100 µm. Thepores are interconnected as shown in Fig.

11 ~ Fig. 13. Fig. 11 shows a kidney scaffold with a volume of ~1400 µm × 820 µm × 700 µm and pore size of ~100 µm. Thepores are interconnected as shown in Fig. 11 (c), where micro-CT imaging (Scanco μCT 80) was used to obtain cross-sectional images and internal pore architectures. Fig. 12 and 13show the cubic and oval scaffold, respectively. The pore sizes are ~85 µm and ~145 µm, respectively, and the volume sizes are ~1280 µm × 980 µm × 1320 µm, and ~1500 µm × 1030 µm × 800 µm, respectively. The scaffolds were fabricated with the thickness of 4 µm. The micro-CT images in Fig. 12 (c) and Fig. 13 (c) also show good pore interconnectivity.However, the cross-sectional areas of the inner pores in Fig. 13 (c) are smaller than the designed pores. This may be due to some residual PPF/DEF solution that was not completely rinsed out of the pores during the rinsing process. Using the PPF/DEF solution, 4 micro-needles were fabricated as shown in Fig. 14.

Using the PPF/DEF solution, 4 micro-needles were fabricated as shown in Fig. 14. Thediameter of the micro-needle’s tip and base are ~20 μm and ~180 μm, respectively, and the height is ~ 800 μm. These images demonstrate the ability of µSL to effectively fabricate micro-needles that could be used as drug delivery devices.(a) (b)(c)Fig. 11 Fabricated kidney scaffold: (a) SEM image, (b) top view of microscope image and measurement, and (c) cross-section of micro-CT image(a) (b) (c)Fig. 12 Fabricated cubic scaffold: (a) SEM image (b) top view of microscope image and measurement, and (c) cross-section of micro-CT image(a) (b) (c)Fig. 13 Fabricated oval scaffold: (a) SEM image (b) top view of microscope image and measurement, and (c) cross-section of micro-CT image(a) (b)Fig.

13 Fabricated oval scaffold: (a) SEM image (b) top view of microscope image and measurement, and (c) cross-section of micro-CT image(a) (b)Fig. 14 Fabricated micro-needles: (a) perspective view, and (b) front viewCONCLUSIONS In this work, 3D micro-scaffolds and micro-needles were fabricated using a previouslydeveloped dynamic mask projection μSL system and the biocompatible polymer PPF. It was confirmed that µSL has the potential to fabricate complex 3D micro-scale structures with interconnecting pores out of a biopolymer. The synthesized PPF was too viscous to be usedin the μSL system, so the viscosity was reduced by adding DEF. Using the PPF/DEF prepolymer and the developed μSL system, a curing experiment was conducted to determine the critical energy and penetration depth for the solution. The ability of the system to fabricate complex 3D micro-scaffolds was demonstrated by showing various scaffold geometries with interconnecting pores.

The ability of the system to fabricate complex 3D micro-scaffolds was demonstrated by showing various scaffold geometries with interconnecting pores. The fabricated micro-scaffolds were observed using SEM and optical microscopy, and the cross-sections were investigated using a micro-CTimaging system. In addition to micro-scaffold fabrication, micro-needles were fabricated todemonstrate the ability of μSL to fabricate biocompatible drug delivery devices. Based on these results, it is concluded that the developed μSL system and the use of PPF show promise for fabricating micro- scaffolds and drug delivery devices with controlled micro-architectures.ACKLOWLEDGEMENTS This work was supported by grant No. R01-2004-000-10507-0 from the Basic Program of the Korea Science & Engineering Foundation, the Mr. and Mrs. MacIntosh Murchison Chair I inEngineering endowment at the University of Texas at El Paso (UTEP), and a research contract from the U.S.

and Mrs. MacIntosh Murchison Chair I inEngineering endowment at the University of Texas at El Paso (UTEP), and a research contract from the U.S. Army Space and Missile Defense Command and the Homeland Protection Institute to the Center for Defense Systems Research at UTEP. The findings and opinions presented in this paper are those of the authors and do not necessarily reflect thoseof the sponsors of this research.REFERENCES [1] Ikuta K., and Kirowatari, K., “Real three dimensional micro fabrication using stereolithography and metal molding,” Proceedings of 6th IEEE Workshop on Micro Electro Mechanical Systems (MEMS’93), New York, NY, 1993, pp.

brication of Curved Ceramic / Polymer Composite Transducers for Ultrasonic Imaging Applications by Fused Deposition of CeramicsG.M. Lous, LA. Cornejo, T.F. McNulty, A. Safari. S.C. DanforthRutgers, the State University of New Jersey, Department of Ceramic and Materials Engineering, 607 Taylor road Piscataway, NJAbstractis a Solid Freeform Fabrication (SFF) technique where a three-dimensional green ceramic object is built layer by layer, starting from a Computer Aided Design (CAD) file of the object. This technique was used to build novel piezoelectric ultrasonic transducers for medical imaging applications. Curved ceramic skeletons for. 2-2 .•(parallel ceramic / epoxy plates) composite transducers were built by FDC. The design's curvature can be tailored in the CAD file. Therefore, the final composite requires very little machining. In the FDC..built green parts, the ceramic plates were 500 Jlm thick and the spacing between the plates was 1270 Jlm.

Therefore, the final composite requires very little machining. In the FDC..built green parts, the ceramic plates were 500 Jlm thick and the spacing between the plates was 1270 Jlm. The FDC green samples were subjected to a slow binder burnout cycle at 550°C for 4 hours, using a heating rate of 8°C per hour, then sintered at 1285°C for 1 hour. Physical characterization of the samples revealed that 95% of the theoretical density was achieved. The ceramic plates shrunk 20% in height as well as in width. The shrinkage was of only 16%.in the direction parallel to the plates. Optical microscopy and SEM were performed on green and sintered samples.

The shrinkage was of only 16%.in the direction parallel to the plates. Optical microscopy and SEM were performed on green and sintered samples. The results of these characterizations are reported in this paper as well as the electromechanical properties of the final composites and of FDC bulk samples.Fused Deposition of Ceramics (FDC), developed at Rutgers University,IntroductionFused Deposition of Ceramics [1] is a Solid Freeform Fabrication technique in which a green c.eramicobjectcan be built layer by layer, due to the deposition of roads by a liquefier moving in the x and y directions. The feedstock for this technique is a .ceramic loaded, 1.78 mm diameter filament, which acts as a piston to push the material out of the nozzle. This technique was used to fabricate the ceramic portion of a piezoelectric ceramic / inactive polymer composite transducer for medical imaging applications.

This technique was used to fabricate the ceramic portion of a piezoelectric ceramic / inactive polymer composite transducer for medical imaging applications. Composites are employed because they combine the property advantages of both polymer and ceramic. transducers and thus have properties superior to the ceramic phase alone [2]. The material of choice was leadzirconate titanate (PZT) forthe ceramic and Spurr epoxy forthe polymer [3]. Ultrasound medical imaging transducers are used both as emitters and as receivers. For optimum efficiency in these two modes, the. volume fraction of ceramic in the composite should be maintained between 20 and 30 volume percent [4]. This range is determined by compromising•to achieve high piezoelectric charge coefficient (d33), moderate acoustic impedance and high thickness coupling coefficient(kt) [5].

This range is determined by compromising•to achieve high piezoelectric charge coefficient (d33), moderate acoustic impedance and high thickness coupling coefficient(kt) [5]. For medical imaging applications, a curvature in the transducer's surface will allow for focusing the beam at a defined distance· away from the transducer.. This. focusing ability leads to a better lateral713resolution, as the beam is smaller in the focal zone and therefore has a higher intensity; this increases the signal to noise ratio [6]. The design flexibility afforded by t1l.e FDC process allows the designer to freely choose the radius of curvature and all other transducer specifications without penalty of cost or manufacturing tim.eassociated with increased complexity of the design.Composite DesignComposites can have different "connectivities" as described by Newnham et al. [7], meaning that each phase can be self interconnected in either one, two or three dimensions.

[7], meaning that each phase can be self interconnected in either one, two or three dimensions. For this project, a 2-2 connectivity was chosen, where both the ceramic and the· polymer are continuous in two directions (x and z), thus forming alternating parallel plates (Figure 1).The minimum green ceramic plate width (w) is dictated by the nozzle diameter used and the firing shrinkage. The volume fraction of ceramic then gives t~espacing (s)between these the walls.•• In order to avoid lateral. resonant waves between the plates (Lamb waves) [8], thickness of the transducer should be··larger than the pitch, which is defined as the sum of the width of one ceramic and one polymer plate. The resonant frequency of the transducer is inversely proportional to the transducer's thickness [6].Itlorder to improve the resolution, it was chosen to use a curved transducer . The radius of curvature should be chosen depending on the application, i.e. the.

The radius of curvature should be chosen depending on the application, i.e. the. distance of the biological organ to image from the transducer. Two radii of curvature were chosen: 10 and 20 cm. Flat transducers (infinite radius of curvature) were also built for comparison.Ceramic o PolymerWidth, wFigure 1: Schematic R.epresentation of a Curved 2-2 Ceramic Composite Built by FDC Technique.Fused Deposition of CeramicsThe first step· in making any structure by. the FDC .technique is ·tofabricate the raw materialf~edstock filarn.ents. The g~nericproc~sing sequence forPZTparticle load~d filatI)ent fabrication was as follows: coatingthe.PZT powder (non spray..dri~d, PZT-5H, TRSCerarnics, Inc., Colleg~ Park, PA) with 1 wt.

The g~nericproc~sing sequence forPZTparticle load~d filatI)ent fabrication was as follows: coatingthe.PZT powder (non spray..dri~d, PZT-5H, TRSCerarnics, Inc., Colleg~ Park, PA) with 1 wt. % stearic acid asa dispersant to lower the syst~msoverall viscosity, compounding of coated powder (to 60 volume fraction of PZTceramic) with bin?~r, granulating the compound~d material into mm sizedp~ll~, and finally extrusion of filaments with close control over their diamet~r. Based on previous results, .stearic acidwaschoseniastll~ dispersant. The binder used isa four component bind~r [9]. Several processing parameters w~re optimized to improve the quality of the filament. Capillary rheometer viscosity measurements714indicate a sufficiently low viscosity (at the actual FDC temperature of 166°C) for properFDC behavior. These filaments were successfully used in the FDC process.The next step was to create a CAD file of the desired transducer.

These filaments were successfully used in the FDC process.The next step was to create a CAD file of the desired transducer. In this project, the ceramic structure had to be curved on the top. However, since SFF techniques are layerwise techniques, the curvature had to be approximated by discrete steps. Figure 2 shows an example of the stepped surface.(a)(b)Figure 2: (a) Computer Aided Design of the 2-2 Composite Structure; (b) Side View of the CAD File Showing the Curvature Approximation by Steps.Quickslicesoftware was th~l1 ut~lizedto slice theCAD.~le and to create the tool path. In order to create the parallel ~ZT phltesforthe2-2 composicte Stntcture, a gap was left between the deposited ceramic roads. For theilrst Parts, the nozzlediaJlleter,road width and spacing between the roads were selected to be 50S Hm,508 J1mand1270Hffir~spectively.The slice thickness was 254 J1m.

For theilrst Parts, the nozzlediaJlleter,road width and spacing between the roads were selected to be 50S Hm,508 J1mand1270Hffir~spectively.The slice thickness was 254 J1m. Later, thanks to th~<i~yelopment of a n~wbinder forrpulation, it was possible to fabricate Parts using a 254J1ffi diarneternozzle; in thes~parts,the road width, spacing and slice thickness were respectively set at330 J.1m,y762 J.1m and 127 J.1m. In each case, the FDC build parameters such as liquefier telllperature, envelope telllperature, flovv rate, etc., were optimized tQjmprovePart quality. The liql1efier and envelope temperature were respectively set at 165· and 50°C. The flow rate was optimizedin order to avoid overflow as well as underflow of material at the point where the nozzle starts depositing.Figure 3 shOWS as-built ceramic Structures for 2-2.composites fabricated ·by FDC with 508i J1m (a)· and 254 J1m (b) diameter nozzles.

It should be noted that the height ofeach of the Parallel PZT plates.isdifferent, which.results in the needed curvature for the focusingyfutlctionof this transducer design.Figure 3: Images ofCurved Green Ceramic Structures Built by the Ditect FDC Technique; (a) Nozzle Diameter = 508J1ITl, (b) Nozzle Diameter::::: 254J.1m.715Post FDC ProcessingThe green PZT structure made by the direct FDC process was subjected to binder burnout (BBO}andsinteringheat treatmt::ntsafter FDC. Since the part hadfine ft::atures,the BBOcycle was modified to maintain part integrity: it was done by heat treating the partin static air atmosphere to 550°C for 4 hours with a heating rate of 10°C per hour. The setup was as follows: the sample was put in an alumina crucible on top of a coarse PZT bed; this bed eased the removal of the binder from the bottom ofthe sample.

The setup was as follows: the sample was put in an alumina crucible on top of a coarse PZT bed; this bed eased the removal of the binder from the bottom ofthe sample. Sintering was performed at 1285°C for one hour in an excess PbOatmosphere (to prevent Pb loss during sintering) in a closed and sealed alumina crucible; the heating rate was 3.5°C per minute. The flow of oxygen inside the furnace was not controlled. This was followed by Corona poling of the ceramic using a voltage of 26 kV for 15 minutes. After aging for 24 hours, the ceramic structure was impregnated in Spurr epoxy polymer. The composites were then electroded, repoled and the electromechanical properties were evaluated.Part CharacterizationBubble&:"-:--.. Small roadFigure 4: Optical Micrographs (and computer enhanced images) of Cross-Sections of Several Sintered FDC Plates made by FDC SFM Process.Optical and scanning electron microscopy was performed on green.

and sintered PZT samples in order to characterize each part's micro- and macro-structure. Defects identified in the first iteration of the process (non-uniform road widths and porosity/bubbles) (Figure 4)· were rectified through process improvements as can be seen from the micrograph shown in Figure 5.Figure 5: Optical micrographs (and computer enhanced image) of cross,,;sections of green PZT plates made by the FDC SFM process, showing that the defects such as small roads and bubbles were eliminated by improved processing.716Table 1 reports the measured values for road widths, slice thickness, pitch and the surface angles. The road width in the green state is whatthe FDC system builds and in the sintered state is after shrinkage during BBO and sintering. The surface angle refers to the angle that the material makes at the intersections of two successive layers.

The surface angle refers to the angle that the material makes at the intersections of two successive layers. The main concern at this point is to decrease both the standard deviations of the values of road thickness, road width and inter...plate thickness (in both the as-built and as-sintered conditions), in order to maximize control over the component and insure that it accurately reproduces what is. specified in the original CAD file.• It is noted that, as with other ceramic materials made by SFF techniques, the green to sintered shrinkage measurements show anisotropy. For these parts, the shrinkage in the direction perpendicular to the roads (20%) was larger than in the direction parallel to the roads (16%). However, this can be .incorporated into the original CAD design and can also be used to an advantage to obtain finer features. Scanning electron microscopy was performed. on a fracture surface(Figure 6).

The sinteredparts were made before processimprovements, the green parts Were mad.e after (See Figures 4 and 5).717Piezo~lectricproperties were measured.ontheflatP~TIpolymercorn,posites an,qonbulk sampl~smadeby FDC. The results are shown in Table· 2.• Th~values obtait).ed areicomparable with those obtained by conventional methods.SampleFlatFDC composite Dice-and-fm composite Bulk sampleVr.PZT (%) 2725100. kt ~ (%) (%) 32 68635471K6277803340tanS0.0230;0850.023«133 (pCIN) 397± 16355 ± 14664 :1::4p (gI~~) 2.882.807.7Z (MRayls) 8.349.7930.54Table 2: Piezoelectric Properties of Composites and Bulk Samples Made by FDC.Summary and conclusionsIn summary, curved PZT composites with 2-2 connectivity have been designed and fabricated by the direct FDC technique. The curved 2-2 composite structure was successfully built by FDC using 508 micron and 254 micron diameter nozzles.

The curved 2-2 composite structure was successfully built by FDC using 508 micron and 254 micron diameter nozzles. Optical microscopy was performed on early samples, showing irregularities in the road widths for the green and sintered samples. Subsequent powder processing improvements have led to significantly better control over defects and road widths, etc. No significant irregularities were observed on recent samples as a result of these processing improvements. The slice thickness showed good uniformity. Scanning electron microscopy showed no delamination and good bonding of the roads. The shrinkage behavior of the FDC samples was shown to be anisotropic, as is the case for other ceramics made by the FDC process.

The shrinkage behavior of the FDC samples was shown to be anisotropic, as is the case for other ceramics made by the FDC process. The piezoelectric properties of the flat composites as well as those of bulk samples made by FDC were found to be similar to those obtained in samples made by conventional methods.AcknowlegmentsThe authors thank: the New Jersey Commission on Science and Technology Multi-lifecycle En'gineering Research Center and the New Jersey Institute of Technology for providing the funding for this project.References[1] S. C. Danforth, A. Safari, Nt. Agarwala, A. Bandyopadhyay, R. van Weeren,N. Langrana, V. Jamalabad and W. Priedeman,~~Solid Freeform Fabrication Methods," U.S. Patent No. 5,738,817, April 14, 1998.[2] T.R. Gururaja, A. Safari, R.E. Newnham and L.E. Cross,ch. 2 in ElectrQnic Cerarp,i9s, edited by L.M. Levinson, Marcel Dekker, Inc., New York, 1987.[3] V.F. Janas and A.

BRICATION OF CUSTOM DYNAMIC PEDORTHOSES FOR CLUBFOOT CORRECTION VIA ADDITIVE-BASED TECHNOLOGIESVito Gervasi, Douglas Cook, Robert Rizza, Ph.D., Sheku Kamara, Milwaukee School of Engineering, Milwaukee, WisconsinXue Cheng Liu, Ph.D, MD Children's Hospital of Wisconsin & Medical College of Wisconsinkeywords: custom orthosis, rapid prototyping, additive manufacturing, laser sintering, foam molding, polypropylene, polyamideABSTRACTAdditive technologies have created many opportunities to improve the quality of life for patients in a range of medical applications. This paper provides an overview of how several of these technologies were leveraged to transform custom pedorthosis designs into physical, end-use, custom pedorthoses for clubfoot patients. The pedorthoses that were produced are currently being tested on the respective patients for their improvement in mobility and degree of clubfoot correction, and will continue through early 2010.

The pedorthoses that were produced are currently being tested on the respective patients for their improvement in mobility and degree of clubfoot correction, and will continue through early 2010. The advantage of this approach is the reduction in labor and the increase in degrees of design freedom available, compared to conventional methods of fabricating orthotic devices. Several new approaches for fabricating custom orthotic devices that were explored, and the related results, are discussed.1. INTRODUCTIONPurpose The purpose of this research was to determine the most-practical means of transforming computer-aided-design (CAD) models of custom clubfoot pedorthoses into functional pedorthoses for testing on patients in a clinical trial. The general pedorthosis design included two mating layers (one of variable thickness), each having specific mechanical properties required to achieve the desired pressure distribution across sole of the patient’s foot.

Precisely producing the variable-thickness regions is a significant challenge for conventional methods that use laminated foams and thermo-formed layers.The goal of this paper is to convey the potential of the fabrication procedure used and lessons learned on this project to the rapid-prototyping and orthotic communities. To that end, the complete procedure and results-to-date are discussed herein.Scope The main challenge of this research was identifying the ideal method for reproducing the somewhat complex, variable-thickness pedorthosis geometry, using appropriate materials.

The targeted goals were: fabricating a custom-fitting pedorthosis to aid in the correction of clubfoot;using a stiff outer shell and low-stiffness inner layer for proper pressure distribution; and, producing a pedorthosis durable enough for daily use.A number of options were considered, including direct additive manufacturing, a hybrid combination of a conventional approach and additive fabrication, and several other more- traditional material-removal approaches. The options explored are all described in the following “Background” section of this paper. Although the approach chosen is not considered the long- term optimal approach, it was the most practical within the timeframe of this project, and can achieve an outcome not obtainable by conventional orthosis-fabrication techniques.

Timing and performance were the primary constraints driving the scope of this project.Background Conventionally, pedorthoses are created by an orthotist using a well-defined, widely accepted (standard) procedure to make a mold from the patient’s leg, requiring several fiberglass and plaster castings and thermal-forming steps. The procedure begins by capturing the foot, ankle and calf in a fiberglass casting applied to the patient’s skin. Upon curing, the “shell” is cut away from the patient1, and re-closed using adhesive tape to form a mold. Next, plaster is poured into the fiberglass-shell mold, resulting in a reproduction of the patient’s lower limb under a no-load condition. Once the plaster is cured, removed from the mold and smoothed, the required foam pads are held in place against the plaster as a heat-softened polypropylene (or polyethylene) sheet is wrapped around the entire setup and crimped shut.

A vacuum is then used to pull the plastic against the plaster and foam pads to form the pedorthosis.For clubfoot patients, these conventional pedorthoses are generally only intended to improve the comfort of the patients. Because the casting was conducted under no-load conditions, correction of clubfoot is not considered.This work supports a collaboration between Dr. Robert Rizza of MSOE and Xue Cheng Liu, Ph.D, MD of Children's Hospital of Wisconsin & Medical College of Wisconsin, that is funded by the U.S. Department of Education under Grant No.

Robert Rizza of MSOE and Xue Cheng Liu, Ph.D, MD of Children's Hospital of Wisconsin & Medical College of Wisconsin, that is funded by the U.S. Department of Education under Grant No. H133G060142, to define a process of mapping a clubfoot patient’s pressure distribution across the sole of their foot as they walk (loaded) and calculate the geometry of the pedorthosis required to correctly re-distribute the loads for a more normal gait.These custom-designed clubfoot-corrective pedorthosis designs have several challenging features that conventional methods of orthosis production cannot accomplish; therefore, the relatively new technologies of 3-D scanning and rapid prototyping (RP) were identified as key enablers for this project.

Unlike traditional material-removal fabrication techniques, such as computer- numerically-controlled (CNC) machining, rapid prototyping is an additive process, starting with an empty build platform, and adding material one layer at a time, on top of the build platform, to create a 3-D object (Figure 1). CAD data is used to define the exterior and interior surfaces of the solid 3-D object, making complex part geometries, as well as complex internal structures, possible and commonplace. RP is a subset of a broader group of layered manufacturing processes titled “Additive Fabrication.”1 Special cutting tools are used to minimize the potential for harming the patient.Figure 1: Conceptual image of Klein-bottle fabrication in layers, exemplifying the capability for directly fabricating complex geometries.Additive manufacturing has made a tremendous impact in changing the methods used to produce many custom personal products, such as hearing aids and dentures.

In a growing number of examples, “traditional manufacturing” methods are no longer competitive with this expanding technology for “growing” one-of-a-kind, end-use parts.AM provides an ideal approach for generating the custom, end-use hard- and soft-layer patterns: each pedorthosis is truly unique; and, the soft layer has regions of variable thickness.2. FABRICATION OPTIONSThe goal of the fabrication portion of this project was to construct a functional dynamic pedorthosis with a hard shell and a soft inner layer, based on designs generated by finite- element-analysis (FEA)-refined CAD models, to accommodate the specific functional requirements. The hard-shell region needed properties to provide strength and support, i.e. serve as the backbone of the pedorthosis, with equivalent or comparable properties to polypropylene or polyethylene currently used.

serve as the backbone of the pedorthosis, with equivalent or comparable properties to polypropylene or polyethylene currently used. The soft inner layer needed to have closed-cell-foam-like properties comparable to the ACOR® ethylene-vinyl-acetate foam product currently used for orthotic padding, with an effective elasticity of 128psi and durometer of 35 Shore A (1).Unfortunately, the aforementioned materials used in conventional orthosis fabrication are not yet available for solid free-form fabrication; therefore, to fabricate the pedorthoses, several approaches were considered, including direct manufacturing, additive-based molding, laser cutting of foam and combinations of several of these approaches.

The soft layer needed to be durable and provide a good recovery after being compressed for a length of time.Direct manufacturing provides an ideal approach since the foam layer, represented by a fine lattice structure whose effective stiffness is lower than the constitutive material (“artificial foam”), and the hard shell could potentially be grown as one integrated part using a process such as 3D Systems’ Selective Laser Sintering® (SLS) or Objet’s PolyJet™ 3-D printing.While 3D Systems does provide a polypropylene material (DuraForm® PP 100) (2), it has only been available for about one year, and only through their Preferred Service Providers (3). The testing required to determine the feasibility of directly manufacturing fine “artificial foam” structures is better done in-house. There is also some concern about the constitution of its blend with respect to prolonged skin contact.

There is also some concern about the constitution of its blend with respect to prolonged skin contact. This material will be pursued further in future efforts.Polyamide (nylon) is an option for SLS (DuraForm® PA) (4); however, there was significant reluctance on the part of several team members in using this nylon material for the immediate application, due to concerns over acceptance by the orthotic community.Regarding PolyJet™ 3D-Printing, the materials are too stiff for the artificial foam (5). Functionally-graded solid materials are quite feasible, and may be used in future applications. Currently, there are no medical-grade materials available from the Polyjet® 3-D-printing technology.Additive-manufacturing-based molding was another option for this project; and, several approaches were available including direct-mold and pattern-based.

The first approach uses additive manufacturing to grow molds derived from the CAD of the desired part, accounting for shrinkage and de-molding challenges.A pattern-based approach can also be used to create a mold, without the need for a respective CAD design. Room-temperature-vulcanizing (RTV) silicone resin can be cast around the positive pattern. By cutting into the cured silicone to remove the pattern, the mold parting surfaces are created. Venting and gating can then be added as needed.For either case the final mold is filled with the desired casting material; and then, upon solidification, the mold is opened to reveal the final part. This approach is desirable because multiple parts can be made from one mold; and, a range of materials are available.

This approach is desirable because multiple parts can be made from one mold; and, a range of materials are available. Undesirable is the added labor of going from a final CAD model of the desired part to a molded part, and the time associated with the added steps.For this project a pattern-based RTV-silicone molding process was used to produce the closed- cell polyurethane inner layer of the pedorthosis. If molds had been grown, the Z-corp 3-D- printing process in plaster would have been the preferred method due to the high speed, low cost, and applicable wax coating (that releases cast polyurethane readily).Laser-cutting of the soft layer was another option that was explored. One prototype and one bi- lateral set of a patient’s orthoses were produced using this method. The soft-layer CAD model was “smashed” using McNeel’s Rhinoceros® 3-D NURBS modeling software (Figure 2, A) resulting in a CAD model that was flattened much like an unwrapped world map (6).

The soft-layer CAD model was “smashed” using McNeel’s Rhinoceros® 3-D NURBS modeling software (Figure 2, A) resulting in a CAD model that was flattened much like an unwrapped world map (6). The pattern was then laser-cut into a closed-cell ACOR® foam (Figure 2, B); and, the resulting foamThese pedorthoses nd adhered to an SLS polyamide hard shell (Figure 3). These pedorthoses cut-out was laid-up and adhered to an SLS producing variable thicknesses was a significant challenge. Additionally, producing variable thicknesses was a significant challenge. were functional; but, producing variable thicknesses was a significant challenge. there was concern that the pedorthosis would deteriorate at there was concern that the pedorthosis would deteriorate at the seams. For these reasons, t method was abandoned for this project. method was abandoned for this project.the seams.

For these reasons, t method was abandoned for this project. method was abandoned for this project.the seams. For these reasons, thisABFigure 2: (A) Rhinoceros® 3-D CAD model of ACOR® foam.D CAD model of a smashed soft-layer pattern and (B) a laser-cut pattern infolded and adhered to a hard SLS polyamide (nylon) shell. The blue foam shell. The blue foam Figure 3: Laser-cut ACOR® foam folded and adhered to support and cushioning of the patient’s is the thinner foam lining; and, the pink foam is thicker for corrective support and cushioning of the patient’s is the thinner foam lining; and, the pink foam is thicker for foot.3. SELECTED APPROACHwas a combination of The approach used to fabricate all of the pedorthoses for the clinical trials was a combination of The approach used to fabricate all of the AM of the hard shell and additiveshell and additive-based molding of the soft layer as follows:A.

Hard Shell: The hard-shell layer was additively manufactured in the Milwaukee School of Engineering’s (MSOE) Rapid Prototyping Center® using 3D-Systems®’ Selective Laser Sintering™ (SLS) with DuraForm® PA (4).Step 1. The CAD model was positioned to align the sole of the orthotic to the build direction(parallel to the build platform).Step 2. The SLS laser settings were appropriately set to a power based on the glaze-testbefore the build to provide a tough, resilient part (Figure 4).Figure 4: Hard shell of one custom pedorthosis produced in the Milwaukee School of Engineering’s Rapid Prototyping Center® on a SLS machine.Step 3. Upon completion of the orthotic, the part was cleared of loose nylon powder. Step 4. Several sharp edges were filed or sanded off to reduce potential for foam wear and discomfort to the patient. These edges were not easily removable in the CAD model due to the complexity and size of the parametric file.B.

These edges were not easily removable in the CAD model due to the complexity and size of the parametric file.B. Soft Inner Layer: The soft layer was fabricated using a RTV-silicone mold to cast thepolyurethane (PU) foam parts. The pattern was produced using MSOE’s stereolithography apparatus (SLA).Step 1. For the silicone and PU used, shrinkage was negligible; therefore, no shrinkage allowance was required in the CAD models. The “.stl” file of the soft layer was positioned for build on the SLA - chosen for its availability, accuracy, smoothness and part density (Figure 5).Figure 5: Light-brown SLA soft-layer pattern, set inside white SLA hard-shell pattern.Step 2. The SLA soft-layer pattern was then positioned in an acrylic mold box with severalcentimeters of clearance on all sides.

The SLA soft-layer pattern was then positioned in an acrylic mold box with severalcentimeters of clearance on all sides. The RTV-silicone (BlueStar Rhodorsil™ 3040A) was poured into the acrylic mold box, until the soft-layer pattern was several centimeters below the surface of the transparent resin (complete immersion). The entire mold box was degassed in a vacuum chamber and allowed to cure overnight, encapsulating the pattern.Step 3. After curing, the mold box was disassembled; and, a razor knife was used to carefullycut through the silicone, along the profile of the pattern, forming a “parting-surface” that allowed for removal of the pattern. The pattern was then removed; and, venting and gating were added as needed. The empty mold (negative pattern) was injected with the desired castable material, in this case a polyurethane (PU) foam, (FlexFoam- iT®-X from SmoothOn™ with a density of 10lbs/ft3 ).

The empty mold (negative pattern) was injected with the desired castable material, in this case a polyurethane (PU) foam, (FlexFoam- iT®-X from SmoothOn™ with a density of 10lbs/ft3 ). Once the foam cured, the mold was re-opened to reveal the solidified part (Figure 6). This approach was most desirable because multiple parts could be made from one mold; and, a range of material densities were available.Step 4. Several fill paths and vents were cut into the silicone to allow for pouring of the castable PU foam into the mold and trapped air to escape, respectively.Step 5. The mold was sprayed with a release agent,closed and positioned for filling with PU.The filling technique that resulted in the greatest success was to pour the mixed PU into the bottom of the mold and quickly, carefully close the mold as the PU begins to expand. The mold was bound with adhesive tape to prevent it from opening when the resin foams.Step 6.

The mold was bound with adhesive tape to prevent it from opening when the resin foams.Step 6. After several hours of curing the PU was removed from the mold and the cast foampart was trimmed of flash and inspected for defects.Figure 6: De-molding of custom, variable-thickness polyurethane foam layer. The mold is RTV silicone.The resulting hard and soft layers were adjusted by the orthotist to fit the patient and finally adhered together using contact adhesive (Figure 7).Figure 7: Mating of custom molded variable-thickness foam insert to stiff additively-manufactured polyamide (nylon) outer shell. The fit is nearly perfect due the accuracy of the fabrication processes involved.4. RESULTS & DISCUSSIONThe chosen approach of additively manufacturing the custom hard shell, and molding the polyurethane-foam insert, resulted in accurate, durable and effective pedorthoses.

RESULTS & DISCUSSIONThe chosen approach of additively manufacturing the custom hard shell, and molding the polyurethane-foam insert, resulted in accurate, durable and effective pedorthoses. The time required for fabrication is 2-3 days, most of which is automated fabrication or material curing.RTV silicone molding was a critical process used to produce the soft layer. This foam-molding process resulted in a good distribution of small cells in the foam. Two parts did develop some unacceptable bubbles; however, they were corrected in successive molds by simply cutting additional vents into the mold to prevent the growth of gas pockets. These pockets can be considered as very large foam cells having a stiffness much lower than the average cell, thereby compromising the performance of the insert.Because of the accuracy of the RP processes, the final parts mated very well.

In fitting the pedorthosis to the patient, the orthotist needed only to detail (sanding, filing, etc.) the hard shell to address “hot spots” (pressure points) before adhering the foam insert. Once fitted, the improvement in gait was dramatic. One patient decided to run down the hallway to show how his pedorthosis was helping him.Five custom clubfoot pedorthoses were fabricated for the five patients (three bi-lateral) in the clinical trial. To date (over six months later), only one soft insert and one hard shell have required replacement. Because each custom insert has a unique, dedicated mold, replacement inserts can be molded in a matter of hours. The clinical testing of the produced pedorthoses will continue through early 2010.Though additive manufacturing was not chosen to produce the entire dynamic pedorthosis, efforts to design and fabricate artificial foam continued in parallel with this research in anticipation of expanding material options.

This development will continue in future work.5. CONCLUSIONSWhile the laser-cutting of foam sheets proved to be a feasible means of producing the soft inner layer, the durability was poor. Due to the time constraints of this project, the most practical fabrication approach was to mold the variable-thickness custom foam insert. Additive manufacturing of the hard shell & additive-based molding of the polyurethane foam insert resulted in accurate, durable and effective pedorthoses that fit well, and could be adjusted as needed. Replacement inserts can be molded in a matter of hours using this silicone-molding approach.Five pedorthoses were successfully fabricated for four patients (one bi-lateral). During the first seven months of use, only one soft insert and one hard shell required replacement.

During the first seven months of use, only one soft insert and one hard shell required replacement. Impact of the dynamic pedorthoses will be measured in late 2009 and early 2010.Future work will further develop “artificial-foam” structures for additive manufacturing of the entire integrated pedorthosis.6. ACKNOWLEDGEMENTSThe authors would like to thank the U.S. Dept. of Education (Grant No. H133G060142) for their support on this project. The authors also extend a thank you to the Medical College of Wisconsin, Children’s Hospital of Wisconsin, and Hanger Prosthetics & Orthotics (particularly J. Kaczmarowski; C.O. of Hanger) for the use of facilities and access to expert support for this project. Finally, the authors thank the efforts of MSOE undergraduate research assistants: Nick Lentz and Dan Neumann for their molding of prototypes.7. REFERENCES1. Acor Orthopaedic, Inc. Acor Brand Microcell Puff. [Online] 2009.

brication of Electrically Conductive, Fluid Impermeable Direct Methanol Fuel Cell (DMFC) Graphite Bipolar Plates by Indirect Selective Laser Sintering (SLS)Kaushik Alayavalli and David L. Bourell,Laboratory for Freeform Fabrication, The University of Texas at Austin, Austin, Texas 78712Reviewed, accepted September 10, 2008Abstract:Graphite bipolar plates are highly desirable due to their properties of high electrical conductivity and low weight but are associated with prohibitive machining costs arising from poor mechanical properties. Bipolar plates made by indirect Selective Laser Sintering (SLS) offer the advantages of complex part production in shorter times at lower cost. Due to the nature of the SLS process, the as-produced (green part) plates are porous and possess low electrical conductivity which can be improved by carbonizing the phenolic resin binder at high temperatures (brown part).

It has been found that the electrical conductivity increases significantly (> 200 S/cm) with a corresponding increase in pyrolyzing temperature which correlates well with literature on the carbonization of phenol formaldehyde resins. The brown parts are subsequently infiltrated with low viscosity (~5 - 10 cps) cyanoacrylate to seal up the open pores, rendering the plates fluid impermeable.1. Introduction:A Direct Methanol Fuel Cell (DMFC) functions by supplying methanol directly to the anode of the fuel cell where it is oxidized to CO2 (Figure 1). The H+ ions are transported across a polymer electrolyte membrane to the cathode where they react with oxygen to form water. A typical DMFC stack is comprised of repeating units of a membrane electrode assembly (catalyst coated electrodes on either side of a polymer electrolyte membrane), diffusion layers, sandwiched between bipolar plates that supply and distribute the fuel, methanol at the anode and oxygen at the cathode.Fig 1.

Schematic of DMFC working principle.Bipolar plates serve the purposes of fuel distribution, current collection, separation of the membrane electrode assemblies and bearing of the clamping force holding the fuel cell stack together. They are the most significant components of a fuel cell stack in terms of weight andcost of production [1]. Graphite has been the standard material of choice for bipolar plates due to its high electrical conductivity and low mass. However, machining fine features such as flow channels and fuel manifolds on graphite plates is expensive since graphite is brittle and has poor mechanical properties [2]. Selective laser sintering (SLS) offers the benefit of producing plates with features of significant complexity at lower cost than machining. Indirect SLS of graphite bipolar plates involves mixing graphite powder with finely milled phenol formaldehyde powder and using a laser to scan individual layers of a .stl file into a solid 3-dimensional object.

Due to the nature of the indirect SLS process, which involves the melting of the phenolic binder by the laser to hold the graphite powder together, the ‘green’ SLS parts produced are porous and require post processing for complete densification. Here, we have demonstrated a cost-effective method of producing bipolar plates by indirect SLS and subsequently carbonizing phenolic resin at > 1000 ºC to obtain electrically conducting porous bipolar plates which are made liquid impermeable by cyanoacrylate infiltration.2. Bipolar plate productionThe process of producing bipolar plates by indirect SLS as shown in Figure 2, involves the production of a green part by the indirect SLS of a mixture of graphite (~ 70 microns) (GS-150E, Graftech Inc) and phenolic (~ 11 microns) (GP-5546, Georgia Pacific) powders in a 70 wt% to 30 wt% ratio, in a Sinterstation 2000 [3]. The porous green part is then heated in a vacuum furnace to carbonize the phenolic binder, thus producing a ‘brown’ part.

The porous green part is then heated in a vacuum furnace to carbonize the phenolic binder, thus producing a ‘brown’ part. Post processing of this brown part includes infiltration with low viscosity cyanoacrylate glue that seals the pores, thereby making the part fluid impermeable, and polishing of the surfaces to expose the underlying electrically conductive graphite matrix. The parameters used for the SLS process are: fill laser power (10 W – 20 W), outline laser power (4 W), powder layer thickness (0.1016 mm), laser scan spacing (0.0762 mm), laser scan speed (1.524 m/s).

Based on Differential Scanning Calorimetry (DSC) results of the phenolic powder, the part bed and feed bins were maintained at 65 ºC and 47 ºC respectively, to ensure that the phenolic powder remained stable without being thermally affected.Powder Preparation & MixingIndirect SLS of Green PartVacuum Furnace CarbonizationLow viscosity polymer infiltrationRoom temperature cure/ PolishingMaterial Properties Characterization and TestingFinished Bipolar PlateFig 2. Flow chart of the experimental process of bipolar plate production by SLSPost processing of the as-produced green parts was aimed at achieving the objectives of high fluid impermeability and electrical conductivity. Initial experiments were aimed at achieving both objectives in one step, viz. infiltrating the pores with a conducting material by electrochemical deposition and vacuum infiltration methods. Cubes of 15 mm side were produced for porosity measurements.

infiltrating the pores with a conducting material by electrochemical deposition and vacuum infiltration methods. Cubes of 15 mm side were produced for porosity measurements. The apparent and actual volumes of these porous structures were obtained by manual measurements and the use of a pycnometer, respectively. Conductivityspecimens of size 40mm by 5mm by 5mm were also made in the same builds and tested using a four probe conductivity setup according to ASTM C611 specifications.3. Results and Discussion3.1 Deposition and Infiltration of electrically conducting polymer:Conducting polymers such as polypyrrole, polyaniline and polythiophene have been reported to show high values of conductivity [4, 5, 6] as well as chemical/thermal stability [7] in the operating conditions of a DMFC.

The objective of the experiments was to synthesize the polymers in situ by:a) Chemical Synthesis: Conductivity and porosity specimens were soaked in a bath of aniline monomer in 1M dopant acidic solution (for conductivity) overnight so as to achieve complete infiltration. While maintaining the bath at 0 ºC, an oxidizing agent such as ammonium persulfate was then slowly added to the bath to initiate polymerization. Upon oxidation, polyaniline separates from the solution and was found to coat the specimens in the solution. The samples were then removed from the solution, dried and then tested for both porosity and conductivity. It was the aim of the experiment to polymerize the monomer present within the porous matrix thereby tackling both issues of conductivity and porosity in one step. Results showed a slight improvement in the conductivity and a similar reduction of porosity in the samples.

Results showed a slight improvement in the conductivity and a similar reduction of porosity in the samples. However, these results did not correspond to minimum standard expectations.b) Electrochemical polymerization: In this technique, the porous graphite specimen is used as the anode in a solution containing the required monomer, suitable solvent (electrolyte) and a doping salt, with a platinum counter electrode. Upon passing a current, electrochemical deposition of polyaniline layers occurs on the porous graphite anode. By initially soaking the graphite plate in the monomer solution, it was hoped to initiate polymerization of the conductive polymer (polyaniline, polythiophene) within the pores. SEM results (Figures 3a and 3b) show considerable deposition of the polymers on the exterior surface. However polymerization in the interior seems limited as a result of surface deposition which prevents access of fresh monomer to the internal pores.Fig 3a.

However polymerization in the interior seems limited as a result of surface deposition which prevents access of fresh monomer to the internal pores.Fig 3a. Electropolymerized Polyaniline on exterior surface Fig 3b. Lack of interior deposition of Polyaniline3.2 Electroless deposition of metalsNickel metal was deposited onto the plate with the aim of infiltration of metal into the porous stru cture thereby increasing the conductivity of the plate. The electroless reduction was performed by placing activated SLSed graphite bars into a bath composed of nickel sulfate (source), sodium citrate (complexing agent), ammonium sulfate (buffering agent) and sodium hypophosphite (reducing agent) [8]. EDS results showed the presence of nickel on the exterior surfaces of the bars but negligible amounts in the interior, contrary to expected results.

EDS results showed the presence of nickel on the exterior surfaces of the bars but negligible amounts in the interior, contrary to expected results. Furthermore, infiltration of metals into the graphite matrix was found to be infeasible, as the metals that would be stable in the environment of the DMFC, according to Pourbaix diagrams, are noble metals such as Gold, Platinum, Palladium, etc., that would be economically infeasible to use [9].3.3 Infiltration using cyanoacrylateCyanoacrylate (ethyl-2-cyanoacrylate) was selected due to its availability in low viscosities cps) so as to achieve complete infiltration of the preforms. It possesses a large operating (~ 5 temperature range (-54 ºC to 121 ºC) and is stable in methanol. Its setting time is extremely rapid (on the order of minutes) and is enhanced by the presence of moisture.

Its setting time is extremely rapid (on the order of minutes) and is enhanced by the presence of moisture. Due to its electrically insulating nature, an additional step for grinding off surface deposition is necessary to restore the electrical properties of the graphite matrix. It can be seen from pycnometric results (Table 1) that the graphite parts made by SLS are > 50% porous. Natural graphite has a theoretical density of 2.26 g/cm3. Upon infiltrating the porous brown part with low viscosity ethyl cyanoacrylate, the pycnometer measured volume (3.5709 cm3) of the part was about 95% of the apparent volume (3.6985 cm3) which shows that complete densification of the part is possible.The cyanoacrylate was infiltrated into the porous specimen by a setup that applied a vacuum on one side of the bipolar plate while the infiltrant was manually dripped onto the other side.

SEM images (Figures 4a and 4b) show charging of the regions containing the infiltrated cyanoacrylate indicating complete penetration into the interior pore structure and the bipolar plate before and after cyanoacrylate infiltration (Figures 4c and 4d). The white discoloration on the plate (Figure 4d) after infiltration is due to the setting of the polymer and is easily removed by polishing the surface. Impermeability to fluids was tested by applying the above infiltration method to liquids such as water and methanol and observing that no seepage from the top of the plate to the bottom occurred.Porosity MeasurementWeight (g) 3 .101 3.2397 3.206 3.287 3.3097 able 1.

It is known that thermal and electrical conductivities of materials are affected by both porosity and non-conducting inclusions. Porosity has been derived by Koh and Fortini [10], to affect conductivity per the equation: J/Jo = (1-ε) / (1+nε2), where J/Jo is the ratio of porous conductivity to fully densified conductivity, ε is the fractional porosity and n is the sensitivity to the presence of pores. Bhatia, et al [11], have shown that the open porosity in phenol formaldehyde resins increases with temperature. Above 800 ºC, the porosity of the carbonized resin drops in a linear fashion with increasing temperature.

Above 800 ºC, the porosity of the carbonized resin drops in a linear fashion with increasing temperature. This data in conjunction with the porosity vs conductivity studies suggests that the conductivity issue can be tackled by reducing the open porosity of phenol formaldehyde resin by carbonizing at high temperatures.The Department of Energy (DOE) has set target parameters fo3.5 Effect of carbonizing temperature on conductivityXRD was performed on the graphite and the phenolic resin before and after pyrolysis at 800 ºC and 1200 ºC (Figures 5a and 5b). The results show that the carbon product of the phenolic nout at 800 ºC is amorphous in nature and therefore may be contributing significantly to the bur drop in electrical conductivity of the part. The XRD result of the carbon produced at 1250 ºC indicates a morphology change that suggests semi-crystallinity.

The XRD result of the carbon produced at 1250 ºC indicates a morphology change that suggests semi-crystallinity. Studies of phenol formaldehyde resin have shown that the electrical resistivity is highly dependent on the pyrolyzation temperature and that there is a change on the order of about 102 between treating at 800 ºC and 1250 ºC [11]. This was verified by measuring the electrical conductivities of selected brown parts and plotting them as a function of the pyrolyzation temperature (Figure 6, tabulated in Table 2). This increase may be due in part to the conversion of amorphous carbon into a semi- crystalline phase and also a reduction in porosity of the phenolic carbonization residue.Fig 5a. XRD at different phenolic burnout temperaturesFig 5b. Unprocessed graphite GS-150EFig 6.

Conclusionsare fluid imper meable and which possess acceptable A process to make bipolar plates that electrical conductivity has been demonstrated. It has been shown that low viscosity ethyl cyanoacrylate polymer infiltrates the specimen completely under vacuum, producing a fully dense part whose electrical conductivity is maintained by grinding the surface to free it of cyanoacrylate. The electrical conductivity of the porous graphite matrix is observed as a function of the burnout temperature, and high values of conductivity (>200 S/cm) have been achieved by varying the same. Research into the change in materials properties (morphology, electrical conductivity and pore structure) of the phenolic resin is underway to better understand the results observed.

Research into the change in materials properties (morphology, electrical conductivity and pore structure) of the phenolic resin is underway to better understand the results observed. Future work also includes the production of bipolar plates based on computational fluid dynamics models aimed at improving fuel distribution across the electrodes and reduction of methanol crossover which is a critical issue in DMFCs.AcknowledgementThis work was supported by the Officeof Naval Research MURI grant No. N00014-07-1-0758.References[1] V. Mehta and J.S.Cooper, Review and Analysis of PEM Fuel Cell Design and Manufacturing, Journal of Power Sources, Vol. 114, 2003, p32 -53.[2] G.O. Mepsted and J.M. Moore, Performace a nd Durability of Bipolar Plate Materials, Handbook of fuel cells – Fundamentals, Technology and applications, John Wiley and Sons Ltd., 2003, vol.

BRICATION OF EMBEDDED HORIZONTAL MICRO-CHANNELSUSING LINE-SCAN STEREOLITHOGRAPHYJae-Won Choi1, Rolando Quintana2, and Ryan B. Wicker1 1W.M. Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, TX 79968 2Department of Management Science and Statistics, The University of Texas at San Antonio, San Antonio, TX 78279Reviewed, accepted September 10, 2008ABSTRACT In an effort to directly and rapidly manufacture micro-fluidic devices with embedded horizontalmicro-channels on the order of tens of microns, a method was developed for using currentcommercially available line-scan stereolithography (SL) technology. The method consisted ofinserting a wire of specified diameter during the build, building around the inserted wire, andremoving the wire once fabricated leaving a channel with a circular cross-sectional geometryequivalent to the wire diameter.

Demonstration of the technique using 31.6 µm, 57.2 µm and 83.5 µm wire was performed using a 3D Systems Viper si2TM SL system and DSM Somos® WaterShedTM resin. By embedding the wire and building around the insert, the down-facing surfaces were supported during fabrication enabling successful and accurate fabrication ofembedded micro-channel geometries. A method for successful fabrication was developed thatinvolved first building an open micro-channel, interrupting the SL process and inserting the wire,and then capping over the wire with multiple layers. After fabricating a part with a micro-wire,the micro-channel was produced by simply pulling the wire out of the part. Scanning electronmicroscope (SEM) images were used to examine and measure the geometries of the fabricatedmicro-channels, and a statistical design of experiments was performed to show that the processwas capable of producing accurate horizontal micro-channels.

It is expected that this processwill enable unique micro-fluidic and other applications of micro-channel fabrication to be pursued using commercial line-scan SL.INTRODUCTION Micro-channels are fabricated by various methods and used in integrated circuits, micro-sensors,and micro total chemical analysis systems (µ-TAS), among others, with the objective of heatexchange or emission, delivery of liquid or gas between fluidic components, diffusion, and so on.The fabrication methods for micro-channels are mostly based on microelectromechanical system (MEMS) technology. One of these methods is bulk micromachining in which a closed micro-channel can be fabricated by bonding an etched channel on a substrate with a regular substrate asa lid [1], by sealing an etched channel with deposition of LPCVD oxide or others [2], or byseeding and growing a diamond on an etched silicon substrate using deep-reactive ion etching(Deep RIE) [3].

In bulk micromachining, additional processes such as bonding or sealing afterproducing the open channel and a clean room are sometimes needed [4], and these result in costincreases and additional process complexity. The size of the micro-channel using thesemethods are from several microns to a few hundred microns and the shape is limited to 2dimensions such as a rectangle because of the nature of the process.Using another MEMS process called surface micromachining, various micro-channels havebeen fabricated [5-8]. This process also leads to additional costs associated with using a cleanroom and manufacturing of a physical mask. With the combination of surface micromachiningand other techniques, a micro-channel fabrication method using SU-8, which is a thick-filmphoto-plastic, was suggested [9]. Chuang et al. [10] fabricated a multi-layer embedded micro-channel by controlling UV dosage and anti-reflection coating on a bottom surface.

Chuang et al. [10] fabricated a multi-layer embedded micro-channel by controlling UV dosage and anti-reflection coating on a bottom surface. Tay et al.[11] suggested direct write of a proton beam on SU-8 to make a micro-channel. While thesemethods are relatively simple, they require an additional bonding process with electronics.Even though the achievable size is as small as a few microns, the cross-section of the fabricatedmicro-channels is again limited to 2 dimensions.In another example, a high aspect ratio vertical micro-channel was fabricated using the LIGAprocess, where LIGA represents a German acronym for a manufacturing process that includeslithography, electro-forming, and molding. Although LIGA can produce highly detailedmicro-features, there are a number of limitations on horizontal channel fabrication [12]. Lee etal. [13] fabricated a micro-channel for an inkjet head by combination of MEMS technology anda Ni electroplating process. Mizuno et al.

Lee etal. [13] fabricated a micro-channel for an inkjet head by combination of MEMS technology anda Ni electroplating process. Mizuno et al. [14] suggested a fabrication method with hot embossand poly(methyl methacrylate) (PMMA) materials. In this process, a silicon wafer was firstetched, a PMMA plate with an open channel was generated by hot emboss, and a closed channelwas fabricated by bonding a PMMA pen channel with a regular PMMA plate. Mizukami et al.[15] fabricated a micro-channel with a smooth surface without a bonding process using the SD method, which is referred to as “Stereolithography with Double controlled surface.” Hawkins et al. fabricated an adhesiveless microfluidic device using solvent assisted thermal welding, where polymer films were used as the material (Mylar sheets), and these sheets were cut and combined according to the fluidic design [16].There are several methods to produce circular cross-section micro-channels using soft materials.

Verma et al. [17] fabricated a polydimethylsiloxane (PDMS) micro-channel by molding PDMS around a nylon thread and pulling out the thread from the mold, leaving a micro- channel. Using this method, Verma et al. fabricated various complex 3D micro-channels withdiameters of ~50 µm to ~250 µm. Golden and Tien [18] fabricated microfluidic hydrogels to be used in tissue engineering applications using molded gelatin as a sacrificial layer, where a micro-channel was as narrow as ~6 µm. Though these methods are promising in 3D complex micro- channel fabrication, they cannot produce solid micro-channels because of the nature of theprocesses and materials. In addition, several studies have demonstrated the fabrication of amicro-channel using a microstereolithography system, which evolved from conventionalstereolithography [19, 20].

In addition, several studies have demonstrated the fabrication of amicro-channel using a microstereolithography system, which evolved from conventionalstereolithography [19, 20]. Using this technology, a biomedical integrated chip (IC) [21], amicro-mixer [22], a micro-ABO blood typing system [23], and artificial human vessels [24] weresuccessfully produced. Using this technology, the micro-channel on a large area over a few cmcannot be easily produced because the systems are focused on producing objects with the rangeof just a few mm.Even though many methods to produce micro-channels have been developed, these methodshave yet to produce long horizontal micro-channels with circular cross-section on the order oftens of microns and contained within a moderately solid and robust package. In our previousresearch [25], vertical and horizontal micro-channel fabrication using commercial line-scanstereolithography (SL) was suggested.

In our previousresearch [25], vertical and horizontal micro-channel fabrication using commercial line-scanstereolithography (SL) was suggested. In line-scan SL, the ability to fabricate vertical micro-channels depends primarily on the spot size of the focused laser, but for horizontal micro-channels, successful fabrication depends on the spot size, cure depth, layer thickness, andpossibly other factors. The cure depth (or the depth of penetration of the laser beam into thephotocurable resin) mostly affects the geometry of the down-facing surface, and it is thischaracteristic of SL that makes horizontal micro-channel fabrication extremely difficult. That is,what typically happens when building the down-facing surface of a micro-channel is the laserwill simply penetrate into the micro-channel completely crosslinking the resin within the channel,and thus entirely eliminating the intended micro-channel.

As a result, SL has not previouslybeen considered as a reliable method for fabricating horizontal embedded micro-channels.In this paper, a method for successful fabrication of horizontal micro-channels usingcommercial line-scan SL has been developed that involves first building an open micro-channel,interrupting the SL process and inserting a micro-wire, and then capping over the wire with multiple layers. After a part with the micro-wire was successfully fabricated, the part was post- cured for a certain time to cure any unreacted resin around the wire. Once post-cured, the micro-channel was obtained by removing the micro-wire from the fabricated part. To demonstrate the ability of this simple and inexpensive process to fabricate these channels, a method for inserting the wire during fabrication was developed that also included the design of a fixture. In the current research, three different wire diameters were used and tested.

In the current research, three different wire diameters were used and tested. SEM images were obtained to examine the geometries of the micro-channels, and the diameters of the micro-channels along the build and scan directions were measured. A statistical design of experiments using measured data was performed to show that the process was capable ofproducing accurate horizontal micro-channels. The following sections contain more details ofthe methods and results obtained from this research.MATERIALS AND METHODSMaterials Micro-wire with nominal diameters of 27.9 µm (0.0011 inch), 53.3 µm (0.0021 inch), and 78.7µm (0.0031 inch) were provided from Phoenix Wire Inc. (South Hero, Vermont). The micro-wires were coated with 2.5 µm (0.0001 inch) thick Teflon using bare wires of 25.4 µm (0.001inch), 50.8 µm (0.002 inch), and 76.2 µm(0.003 inch) diameters, respectively.

The micro-wires were coated with 2.5 µm (0.0001 inch) thick Teflon using bare wires of 25.4 µm (0.001inch), 50.8 µm (0.002 inch), and 76.2 µm(0.003 inch) diameters, respectively. The nominaldimensions were provided from the manufacturer, and the actual individual wire dimensionswere measured as will be discussed later. The three coated micro-wire sizes will be referred toas mw1, mw2, and mw3 for convenience. Micro-channel specimens were fabricated using a Viper si2TM SL system (3D Systems Inc.) and DSM Somos® WaterShedTM 11120 resin.Fixture model To more easily fabricate micro-channel specimens and handle the wire, a fixture was designed asshown in Fig. 1. The fixture included aligning and tying posts on each side of the specimen,where each post was used to secure the wire on each end of the specimen. Each end of the wirewas wrapped around an aligning post and secured by tying to a tying post.

Each end of the wirewas wrapped around an aligning post and secured by tying to a tying post. The fixture wasdesigned for seven specimens where each specimen was built on top of a base plate. Eachspecimen had guidance grooves built into the fixture for wire alignment. The tying posts werealso capped with ring fasteners that were used for wire tensioning.Fig. 1 Wire embedding fixtureSpecimen model The specimens consisted of 4 types of layers as shown in Fig. 2. Fig. 2 (a) and (b) were usedfor base layers with the thickness of 2.2352 mm, Fig. 2 (c) was used for the open channel withthe width of 0.1016 mm and the thickness of 0.1524 mm (3 layers), and Fig. 2 (d) was used forcapping layers with the thickness of 0.254 mm (5 layers), where the stacking layer thickness wasused as 50.8 µm (0.002 inch) in high resolution mode in the machine. Fig.

2 (d) was used forcapping layers with the thickness of 0.254 mm (5 layers), where the stacking layer thickness wasused as 50.8 µm (0.002 inch) in high resolution mode in the machine. Fig. 3 shows the 3Dmodel of the specimens with the size of 10 mm × 12 mm × 2.489 mm, where the wedge groovewas used to assist with breaking the specimens into two parts in order to inspect the inside of thefabricated micro-channel.(a) (b) (c) (d)Fig. 2 Sliced layers for the specimens: (a) bottom layers, (b) intermediate layers, (c) openchannel layers, and (d) capping layersFig. 3 Specimen modelMicro-channel fabrication After fabricating the fixture, a Mylar sheet was put on top of the base plate to act as the specimen build platform so that the specimens could be built without base support and the fixture could bereused as needed. The micro-channel specimen was built on the Mylar sheet by fabricating anopen channel in the specimen (Fig.

The micro-channel specimen was built on the Mylar sheet by fabricating anopen channel in the specimen (Fig. 4 (a)), inserting the micro-wire within the open channel (Fig.4 (b)) while interrupting the machine, capping over the micro-wire with the designed cappinglayers (Fig. 4 (c)), removing the specimen from the fixture after cutting the ends of the micro-wires (releasing the wires from the posts), rinsing the sample with isopropyl alcohol (IPA), post-curing for ~30 min on each side, and pulling the micro-wire out of the specimen (Fig. 4 (d)).For precisely embedding micro-wires on the open channel, one of the ends of micro-wire wastied at the tying post, and it was aligned using the aligning post and the wedge guidance groove.To maintain wire tension during the insertion process, the other end of the micro-wire was tied atthe other tying post as the wire was manually inserted into the open channel.

Once inserted andcapped, each specimen was post-cured in order to completely solidify uncured resin around theembedded micro-wire. To remove the wire, one end of micro-wire was picked up with a tool(needle nose pliers), and the wire was manually and slowly removed from the specimen. Thisprocess is simple and straightforward, although considerable development was required todevelop the final process described here. Much of the preliminary wire embedding researchfrom which the current work is based can be found in Ranade [27].(a) (b) (c) (d)Fig. 4 The process for micro-channel fabrication: (a) fabrication of open channel, (b) embeddingthe micro-wire, (c) capping, and (d) pulling the micro-wireMetrology To verify that the fabricated micro-channel exists with the circular shape in the specimen, the inside of the specimen has to be inspected. The specimen was designed to be easily broken by the wedge groove as shown in Fig. 3.

The specimen was designed to be easily broken by the wedge groove as shown in Fig. 3. After removing the micro-wire, the specimen was split into two parts as shown in Fig. 5 (a). To inspect the micro-channel, a scanning electron microscope (SEM) was used and each end of the split part was inspected as shown in Fig. 5 (b). The micro-channel was measured along the build direction and scan direction as shown in Fig. 5 (b) and these measurements were recorded. For statistical analysis, four builds consisting of five specimens each were fabricated for each of three different size micro-wires. All specimens in each build were randomly inspected using the SEM.(a) (b)Fig. 5 Broken specimen for inspection: (a) broken specimen, and (b) specimen setting forinspectionMICRO-CHANNEL FABRICATION AND INSPECTIONMicro-channel fabrication The fabricated fixture and specimen are shown in Fig. 6 (a), and the embedded micro-wire withthe fabricated specimen is shown in Fig. 6 (b).

6 (a), and the embedded micro-wire withthe fabricated specimen is shown in Fig. 6 (b). After embedding the micro-wire, capping overthe micro-wire, and cutting the micro-wire, the specimens on the Mylar sheet were removed asshown in Fig. 7. These specimens were rinsed with the IPA, and post-cured for ~30 min oneach side. Fig. 8 shows post-cured specimens with the micro-wires, and the micro-wire waspulled out of the specimen as shown in Fig. 9 (a). To measure the inside of the micro-channelthe post-cured specimens were split into two parts as shown in Fig. 9 (b).(a) (b) Fig. 6 Fixture and specimen: (a) fabricated fixture and specimen in the machine, (b) magnifiedimage of the fixture and specimenFig. 7 Fabricated specimens on the Mylar sheetFig. 8 Specimens after post-curing(a) (b)Fig. 9 Specimens after pulling the micro-wire: (a) specimens, (b) broken specimensMicro-wire measurement The diameters of the bare wires (Fig. 10) and coated wires (Fig.

9 Specimens after pulling the micro-wire: (a) specimens, (b) broken specimensMicro-wire measurement The diameters of the bare wires (Fig. 10) and coated wires (Fig. 11) were measured for comparison with the reported nominal diameters. The measured diameters of the bare wires were 28.2 µm, 54.2 µm, and 78.1 µm, respectively, which are larger than the nominal wire diameters of 25.4 µm, 50.8 µm, and 76.2 µm, respectively. The coated wires (Fig. 11) were also larger than their reported values: 31.6 µm for the 27.9 µm (0.0011 inch) wire, 57.2 µm for the 53.3 µm (0.0021 inch) wire, and 83.5 µm for the 78.7 µm (0.0031 inch) wire. Based on the measured values, the approximate Teflon coating thicknesses are 3.5 µm, 3.1 µm, and 5.4 µm,respectively.(a) (b) (c)Fig. 10 SEM images of bare micro-wires: (a) 25.4 µm, (b) 50.8 µm, and (c) 76.2 µm wire(a) (b) (c)Fig.

10 SEM images of bare micro-wires: (a) 25.4 µm, (b) 50.8 µm, and (c) 76.2 µm wire(a) (b) (c)Fig. 11 SEM images of Teflon coated micro-wires: (a) mw1 (b) mw2, and (c) mw3Micro-channel metrology Using the method to produce the micro-channel described above, circular micro-channels weresuccessfully fabricated as shown in Fig. 12 for wires of three different diameters (31.6 µm, 57.2µm, and 83.5 µm wire). To measure the diameters of the micro-channel, the inside of thecircular channel was measured along the scan direction (horizontal direction), and build direction(vertical direction). The results demonstrate that the method described here can be used tofabricate precise micro-channels with circular cross-section.(a) (b) (c) Fig. 12 SEM images of micro-channel: (a) mw1, (b) mw2, and (c) mw3STATISTICAL ANALYSIS The following describes and provides rationale for the experimental design and presents theexperimental results and concomitant statistical analysis.

The key assumptions underlying theadequacy of the statistical models are then presented. The completely randomized factorialexperiment on the fabricated micro channels involved the study of the effects of four factors(Build, Specimen, Diameter or initial mean wire diameter, and Orientation – X and Y measuredchannel diameter) on the difference between the Diameter and the measured diameter of thefabricated micro channel (herein referred to as Diff). This preliminary investigation wasfocused on determining what effects the factors have on the diameter of the fabricated microchannel.In each complete replication of the experiment, all possible combinations of the levels of thefactors (4 for Build, 5 for Specimen, 3 for Diameter and 2 for Orientation) were investigated.All 60 parts were made from the material batch from the same manufacturer.

Note that ε, the random error component, has a single level corresponding to a single replicate per cell.Experimental results Table 1 shows the individual specimen measurements for all of the samples used in the statisticalanalysis. Several observations can be made based upon the experimental results contained intable 1. First, the overall average micro-channel diameters for each of the three diameters areconsistently smaller than the diameter of the embedded wires themselves (except for themeasured Y diameter for the 57.2 µm wire). This is a curious result requiring additionalinvestigation, although the maximum diameter difference is no more than 2.7 µm, which mayresult from deformation of the Teflon coating on the wire during fabrication. There are severalother possibilities for this difference and we are currently investigation this issue.

Dev.29.58731.15550.62302568 0.8297317355.946558.96150.88843275 1.3038223680.81482.76551.41532823 1.49727578Table 2 shows the mean Diff for each level of the factors. It also shows the standard error ofeach mean, which is a measure of its sampling variability. The rightmost two columns show 95.0% confidence intervals for each of the means. Fig. 13 depicts these results graphically.Note that only the Diameter factor shows clear differences between all of its levels. This keyfinding is determined analytically in the next section.Multifactor analysis A multifactor Analysis of Variance (ANOVA) for the dependent variable Diff by levels of theBuild, Specimen, Diameter and Orientation factors is first presented.

Various tests and graphsto determine which factors have a statistically significant effect on the dependent variables wereperformed, as well as tests for significant interactions amongst the factors.The multifactor ANOVA, summarized in Table 3, decomposes the variability of thedependent variable into contributions due to the various factors [26]. All F-ratios are based onthe residual mean square error.

Further, Type III sums of squares was utilized since it allows forthe contribution of each factor to be measured after having removed the effects of all otherfactors.Table 2 Least squares means for Diff with 95.0 percent confidence intervalsLevelCountMeanStandard Error Lower Limit Upper LimitGrand Mean1200.895Build 1300.1099330.2166531.528730.669934Build 2300.7216670.2166531.151070.292267Build 3301.115670.2166531.545070.686267Build 4300.6433330.2166531.072730.213934Diameter 31.6401.228750.1876271.600620.856879Diameter 57.2400.2540.1876270.1178710.625871Diameter 83.5401.710250.1876272.082121.33838Orientation 1601.984170.1531972.28781.68054Orientation 2600.1941670.1531970.1094650.497798Specimen 1241.156670.2422251.636750.676583Specimen 2240.7704170.2422251.25050.290333Specimen 3241.166670.2422251.646750.686583Specimen 4240.9258330.2422251.405920.44575Specimen 5240.4554170.2422250.93550.0246668Vertical scale = -2.7 to 2.3BuildDiameterLevel 1 2 3 4 5OrientationSpecimenBuildDiameter Orientation SpecimenFig.

13 Factor means plot for DiffTable 3 Analysis of variance for Diff – Type III sums of squaresSourceSum of SquaresDf Mean SquareF-RatioP-ValueA: Build5.5147931.838261.310.2764MainB: Diameter83.8489241.924529.770.0000EffectsC: Orientation142.3541142.354101.090.0000D: Specimen8.4474642.111861.500.2073Residual153.4891091.40815Total (Corrected)393.654119All F-ratios are based on the residual mean square error.The above analysis indicates that Diameter and Orientation, with p-values ~0, are the onlyfactors that have a significant effect on Diff. Tables 4 and 5 apply a multiple comparisonprocedure to determine which means are significantly different from which others for Diff by the significant factors Diameter and Orientation. The bottom half of the output shows the estimated difference between each pair of means.

The bottom half of the output shows the estimated difference between each pair of means. Within each column, the levels containing X's form a group of means within which there are no statistically significant differences. The method used to discriminate among the means is Fisher's least significant difference (LSD) procedure.

With this method, there is a 5.0% risk of calling each pair of means significantly different when the actual difference equals 0 [26].Table 4 Multiple range tests for Diff by DiameterMethod: 95.0 percent LSDDiameterCountLS MeanLS SigmeHomogeneousGroups83.5401.710250.187627X31.6401.228750.187627X57.2400.2540.187627XContrastDifference± Limits31.6-57.2*-1.482750.52590531.6-83.50.48150.52590557.2-83.51.964250.525905denotes a statistically significant difference.Table 5 Multiple range tests for Diff by OrientationMethod: 95.0 percent LSDOrientationCountLS MeanLS SigmaHomogeneousGroups1601.984170.153197X2600.1941670.153197XContrastDifference± Limits1-2*-2.178330.4294denotes a statistically significant difference.The following summarizes these findings:Since the p-values for Build and Specimen were greater than .05, we conclude that these factors have no effect on Diff at the 95% confidence level.

However, differences between 2 contrasts in Specimen with respect to Diff were found, shown in Fig. 14.There are statistically significant differences between the mean values of Diff from one level of Diameter and Orientation (shown by an asterisk next to the pairs in the preceding tables), respectively, to another at a level of confidence of 95%, as depicted in Fig. 15 and 16.Means and 95.0 Percent LSD Intervals0.10.4f f i0.7D11.31.6123 Specimen45Fig. 14 means and 95 % LSD intervals between levels of Specimen and DiffMeans and 95.0 Percent LSD Intervals10.50f f iD0.511.5231.657.2 Diameter83.5Fig. 15 Means and 95 % LSD intervals between levels of Diameter and DiffMeans and 95.0 Percent LSD Intervals0.80.30.2f f iD0.71.21.72.22 1 OrientationFig.

15 Means and 95 % LSD intervals between levels of Diameter and DiffMeans and 95.0 Percent LSD Intervals0.80.30.2f f iD0.71.21.72.22 1 OrientationFig. 16 and 95 % LSD intervals between levels of Orientation and DiffTable 6 summarizes the significance of factor levels versus Diff at a level of confidence of95.0 %. Recall that Diameter and Orientation were the only factors that had a statisticallysignificant effect on the dependent variable Diff at the 95.0 % level of confidence. As expected,the diameter of the embedded wire has a statistically significant effect on the resulting channeldiameter. However, what is less obvious is the significance of Orientation on the micro-channel diameter. As noted in the discussion for Table 1, the measured channel diameter in the builddirection (Y Diameter) is consistently greater than the diameter in the scan direction (XDiameter).

As noted in the discussion for Table 1, the measured channel diameter in the builddirection (Y Diameter) is consistently greater than the diameter in the scan direction (XDiameter). The maximum difference between the X and Y diameters is 4.6 µm but is moretypically on the order of 1 µm. The explanation for this result is requiring additionalinvestigation.Table 6 Significance summary of Diff by factor at the 95 % level of confidenceFactorBuildSpecimenDiameter OrientationLevel 1 2 3 4 1 2 3 4 51 2 312Build1234Specimen12345Diameter123Orientation12Finally, Table 7 shows the analysis of variance table which divides the variance of Diff into 4components, one for each factor, in order to estimate the amount of variability contributed byeach of the factors (variance components). Each factor after the first is nested in the one above.In this case, the factor contributing the most variance is Orientation.

Each factor after the first is nested in the one above.In this case, the factor contributing the most variance is Orientation. Its contribution represents71.22% of the total variation in Diff. The effect of Orientation is seemingly very important andhas become a focus of current investigations.Table 7 Analysis of Variance for DiffSourceSum ofSquaresDfMean Square Var.

Comp.PercentTotal(Corrected)393.654119Build5.5147931.838260.00.00Diameter94.7008811.83760.00.00Orientation183.6091215.30072.8313371.22Specimen109.83961.144061.1440628.78Model adequacy In all models generated, no indications were present to assume that the fundamental conditionsfor the models were violated, based on the following analyses:The normal probability plots do not reveal violations of the normal and identicallydistributed (NID) error term.The residuals are structureless; that is, the plot of residuals versus fitted values of thedependent variables did not reveal any obvious patterns.Standardized residuals were approximately normal with mean = 0 and unit variance – no residual was greater than 2 standard deviations, and thus outliers did not seriously distort the ANOVA.The residual plots showed no significant tendencies, and thus no correlation betweenresiduals.No 2-factor interactions in the models were present.

• The assumption that the variance of the observations did not change significantly as the magnitude of the observation changed is upheld utilizing Cochran’s C, Hartley’s and Levene’s tests [26]. It should be noted that the measuring instruments were calibratedbefore each test.CONCLUSIONS A method for fabricating circular micro-channels with diameters of a few tens microns wasdeveloped using commercial line-scan stereolithography, and this method was demonstrated using a 3D Systems Viper si2TM SL system and DSM Somos® WaterShedTM resin. The method consisted of inserting a wire of specified diameter during the build, building around the insertedwire, and removing the wire once fabricated leaving a channel with a circular cross-sectionalgeometry equivalent to the wire diameter. Teflon coated micro-wires of diameters 31.6 µm,57.2 µm and 83.5 µm were used and a fixture was designed to provide for easy wire handling and fabrication of multiple specimens.

Teflon coated micro-wires of diameters 31.6 µm,57.2 µm and 83.5 µm were used and a fixture was designed to provide for easy wire handling and fabrication of multiple specimens. The fabricated specimens were split into two parts to inspect the inside of the micro-channels, and the micro-channel diameters were measured using an SEM. The measured data showed that the micro-wire was successfully removed from the channel, leaving high quality micro-channels with diameters similar to the embedded wire. A statistical analysis was conducted, and it is concluded that the suggested process is capable of producing horizontal circular micro-channels with diameters that vary with the embedded wire diameter. However, the statistical analysis also found that the diameter of the micro-channel was statistically different depending on whether it was measured in the scan direction or builddirection.

However, the statistical analysis also found that the diameter of the micro-channel was statistically different depending on whether it was measured in the scan direction or builddirection. This result was unexpected and we are currently investigating possible sources forthese differences. In conclusion, because this process is simple and fast, we believe line-scanstereolithography can be used to produce a wide variety of unique micro-fluidic and other micro-channel devices.ACKNOWLEDGMENTS The facilities within the W.M. Keck Center for 3D Innovation (Keck Center) used here containequipment purchased through Grant Number 11804 from the W.M. Keck Foundation, a facultySTARS Award and two equipment grants from the University of Texas System, and twoequipment grants from Sandia National Laboratories.

Keck Foundation, a facultySTARS Award and two equipment grants from the University of Texas System, and twoequipment grants from Sandia National Laboratories. Frank Medina, Keck Center Manager,designed the original wire embedding fixture from which much of this work was based, trainedJWC on how to embed wires during stereolithography fabrication, and assisted in various waysduring this research, and his efforts are appreciated. The SEM images were acquired using anSEM funded through NSF grant DMR-0521650. This material is based in part upon worksupported through the Mr. and Mrs. MacIntosh Murchison Chair I in Engineering, researchcontract 504004 from Sandia National Laboratories in the Laboratory Directed Research andDevelopment (LDRD) program, and a research contract from the U.S. Army Space and MissileDefense Command and the Homeland Protection Institute to the UTEP Center for DefenseSystems Research.

Army Space and MissileDefense Command and the Homeland Protection Institute to the UTEP Center for DefenseSystems Research. Sandia National Laboratories is a multi-program laboratory operated bySandia Corporation, a Lockheed Martin Company, for the United States Department of Energy’sNational Nuclear Security Administration under contract DE-AC04-94AL85000. The findingsand opinions presented in this paper are those of the authors and do not necessarily reflect thoseof the sponsors of this research.REFERENCES [1] Nakanishi H., Nishimoto T., Nakamura N., Nagamachi S., Arai A., Iwata Y., and Mito Y.,“Fabrication of electrophoresis devices on quartz and glass substrates using a bonding with HFsolution,” in Proc. IEEE MEMS’97, Nagoya, Japan, Jan. 26-30, 1997, pp. 299~304. [2] Kaplan W., Elderstig H., and Vieider C., “A novel fabrication method of capillary tubes on quartz for chemical analysis application,” in Proc. IEEE MEMS’94, Int. Conf. MEMS, Oiso, Japan, Jan.

Conf. on Solid-State Sensors and Actuators, and Eurosensors IX,Stockholm, Sweden, June 25-29, 1995, pp. 321~324.[6] Liu C.W., Gau C., and Dai B.T., “Design and fabrication development of a micro flow heatedchannel with measurements of the inside micro-scale flow and heat transfer process,” Biosensorsand Bioelectronics, Vol. 20, 2004, pp. 91~101.[7] Liu C.W., Gau C., Ko H.S., Yang C.S., and Dai B.T., “Fabrication challenges for acomplicated micro-flow channel system at low temperature process,” Sensors and Actuators A130-131, 2006, pp. 575~582.[8] Ko H.S., Liu C.W., Gau C., and Yang C.S., “Fabrication and design of a heat transfer micro-channel system by a low temperature MEMS technique,” Journal of Micromechanics andMicroengineering, Vol. 17, 2007, pp. 983~993.[9] Guerin L.J., Bossel M., Demierre M., Calmes S., and Renaud P., “Simple and low costfabrication of embedded micro-channels by using a newthick-film photoplastic,”Transducers ’97 Int. Conf.

Conf. on Solid-State Sensors and Actuators, Chicago, June 16-19, 1997, pp.1419~1421.[10] Chuang Y.J., Tseng F.G., Cheng J.H., and Lin W.K., “A novel fabrication method ofembedded micro-channels by using SU-8 thick-film photoresists,” Sensors and Actuators A, Vol.103, 2003, pp. 64~69.[11] Tay F.E.H., Kan J.A., Watt F., and Choong W.O., “A novel micro-machining method for thefabrication of thick-film SU-8 embedded micro-channels,” Journal of MicromechanicsMicroengineering, Vol. 11, 2001, pp. 27~32.[12] Liu R.H., Vasile M.J., Goettert J., and Beebe D.J., “Investigation of the LIGA process tofabricate microchannel plates,” Transducers’97 Int. Conf. on Solid-State Sensors and Actuators,Chicago, June 16-19, 1997, pp.

brication of In-situ SiC/C Thermocouples by Selective Area Laser DepositionLianchao Sun, Kevin J. Jakubenas, James E. Crocker, Shay Harrison, Leon L. Shaw, Harris L. Marcus Institute ofMaterials Science The University of Connecticut, U-136 97 North Eagleville Road Storrs, Connecticut 06269-3136AbstractWith the intrinsic nature to process relatively small features, selective area laser deposition (SALD) is a potential technique to fabricate complex shaped macro-components with in-situ high-resolution micro-devices. In this study, SALD was used to deposit in-situ silicon carbide/carbon (SiC/C) thermocouples on alumina and silicon carbide substrates with a C02 laser. Tetramethylsilane (TMS) and acetylene (C2H2) were chosen as precursors for deposition of the silicon carbide and carbon lines respectively. The electromotive force (emf) of the deposited thermocouple was measured and found to respond sensitively to temperature variations from room temperature to 800°C.

The electromotive force (emf) of the deposited thermocouple was measured and found to respond sensitively to temperature variations from room temperature to 800°C. The effect of the deposition parameters on the product morphology was also investigated.1 IntroductionAmong the many advantages of Solid Freeform Fabrication (SFF) technology is the possibility of building geometrically complex shapes, tailoring functionally graded materials or composites spatially, and embedding electromechanical devices in situ within parts during the fabrication process. An approach to these goals is to combine two SFF methods, Selective Area Laser Deposition (SALD) and the closely related SALD-Vapor Infiltration (SALDVIil S - ]. SALDVI is used to build the bulk structure of the part while SALD is used to embed the sensors. Many kinds of sensors, such as thermistors, thermocouples, strain gauges, etc., could be formed based on this approach.

Many kinds of sensors, such as thermistors, thermocouples, strain gauges, etc., could be formed based on this approach. Since SALD and SALDVI are completely compatible gas phase approaches, whole parts can be fabricated in a single integrated SALD/SALDVI process system As a first attempt using the SALD/SALDVI technique, thermocouples have been embedded within a silicon carbide substrate. This paper discusses the SALD formation ofthe thermocouple; the SALDVI substrate is described elsewhere in these proceedings [6].Forming a good thermocouple requires consideration of several factors including compatibility with the substrate, formation of a strongly bonded hot end junction, connections from the cold end to readouts, and the physical stability. To avoid thermal residual stresses, insulation, and differences in the thermal expansion coefficient between the substrate, thermocouple lines must be as small as possible.

To avoid thermal residual stresses, insulation, and differences in the thermal expansion coefficient between the substrate, thermocouple lines must be as small as possible. Obviously, embedding SiC/C thermocouples into a SiC substrate has the advantage ofthe low thermal expansion mismatch, which will help to maintain continuity ofthe thermocouple lines during heat/cool cycles.The first SiC/C thermocouple was developed to measure the temperatures of molten steel by Fitterer in 1933[7]. This thermocouple was formed through simply putting SiC and graphite rods together. An extremely high electromotive force constant, up to 300fJ,VfC, was reported. But, this thermocouple is not stable in both liquid metals and gas environment. This greatly these hindered the further application of SiC/C thermocouple.

But, this thermocouple is not stable in both liquid metals and gas environment. This greatly these hindered the further application of SiC/C thermocouple. With the SALD technique, problems could be avoided by enclosing SiC/C thermocouples within a SiC substrate.4812 Experimental procedures 2.1 Procedures ofthe device fabricationAn overall process schematic to make the embedded thermocouple device is shown in Figure 1. The substrate is fIrst built layer by layer with SALDVI. An electrically insulating layer is then put on this substrate. With different gas precursors, the thermocouple lines are deposited on the insulation layer by SALD. The fInal steps of building device are repeating the electrically insulating layers and continuing to build the remainder of the substrate with SALDVI. Single point or multiple-point thermocouples could be made in one layer through this process. In addition, multiple thermocouples could also be built within different layers.

Single point or multiple-point thermocouples could be made in one layer through this process. In addition, multiple thermocouples could also be built within different layers. This will be useful when the multi-point temperature control is needed.Substrate made by SALDVIComplete parts by SALDVI~ Insulating Layer~,,~ Metal Leads~~~l SiC Substrate~ Powder Holder& PistonFigure 1 Flowchart of fabricating in-situ device embedded in macro-component4822.2 Deposition of silicon carbide line, carbon line and insulation layerSALD was carried out using a SFF system which is comprised of a reaction chamber, powder delivery system, 2-D laser beam scanning control stage, gas precursor supply system, observation and recording system, closed-loop temperature control system, and a C02 laser with a maximum output of50W in cw mode and a wavelength of lO.64J.lIll. To get a theoretical spot size of about 500/-lm, a 254mm focal lens was used. Maximum power density is about 2.

To get a theoretical spot size of about 500/-lm, a 254mm focal lens was used. Maximum power density is about 2 . 255W/mmTetramethylsilane Si(CH3)4 (TMS) was chosen as the precursor for deposition of SiC because it produces silicon carbide free from halogens and the by-products of pyrolysis reaction are neither corrosive nor contaminant.For carbon deposition, several precursors, acetylene, ethylene, and methane, wereevaluated. The overall reactions are as followings:C2H 2(g) ¢:> 2C(s) + H 2(g) C2 H 4(g) ¢:> 2C(s) + 2H2(g) CH4(g) ¢:> C(s) +2H2(g)Insulation layers were deposited with a gas mixture of TMS and NH3• Preliminary thermocouple depositions were carried out on alumina and silicon carbide substrates, which were made with conventional sintering method.

Additional experiments were conducted on SALDVI silicon carbide substrates.2.3 Measurement ofelectromotive force (emf)The emf of the thermocouple was measured by connecting the cold end directly to the measurement unit through copper wires. Strawberry tree software was applied to collect the emf signals. The laser beam was used to heat the hot end of the thermocouple. A standard K type thermocouple was also put on the hot end as an independent temperature sensor to monitor the temperature change ofthe hot end. The measurement setup is shown in Figure 2.Alumina SheetCopper SheetJCopper Wire\.. Copper ClampSubstrateInsulation SheetKTCFigure Z Setup for measuring emf of the SiC/C thermocouple2.4 Characteristics ofthe deposited productSALD thermocouple line morphologies were analyzed with scanning electron microscopy (SEM).

The thickness profile of the lines perpendicular to the scan direction was examined with a contact profilometer.483Copper Wire3 Results and discussion 3.1 Silicon carbide linesA series of SiC lines were deposited on the sintered alumina, the sintered SiC substrates and SALDVI SiC substrates. It was found that line profiles were affected by the deposition conditions oftemperature and gas pressure. Generally, with the increase oftemperature and TMS gas pressure, the line shape changes from a sharp peak to a flat top to a coarse, rough top surface. In addition, both width and thickness increase with temperature and gas pressure at the range of the present research.

In addition, both width and thickness increase with temperature and gas pressure at the range of the present research. These results are listed in Table 1; the typical profiles of deposited lines are shown in Figure 3.Table 1 Width and thickness of deposited silicon carbide lines on the alumina substrate ( scanning speed: 0.5mm1sec; 8 passes)TMSgas pressureLine width And Thickness (J1m)700900SALD Temperature (lC)10001100125620 TorrWidtb/5106207561040Thickness/13.928.842.158.340 TorrWidth/3905708101195Thickness/31.9114.1142.9143.860 TorrWidth375410/830/Thickness16.568.8/147.8/~'-'.31»,1100 C.. , / ; ~-~: ,·_··J"O")J··········-··_··r~··_··_,-_·····'l~··_··_,·~··• ' ••••• :....... ~.\.".

, / ; ~-~: ,·_··J"O")J··········-··_··r~··_··_,-_·····'l~··_··_,·~··• ' ••••• :....... ~.\.". r"L..-~ Width (J11l1)Figure 3 Typical profile cbanges of the deposited SiC lines (Pntr20 Torr; Scanning speed: 0.5mm1sec; C02 laser; Alumina substrate; 8 passes)4843.2 Carbon depositionThe deposition rate of carbon is relatively slow compared to that of silicon carbide especially when methane or ethylene was used. This is consistent with other research results [8]. After the preliminary experiments, acetylene was chosen for deposition of carbon lines. Three kinds of morphologies, cotton-like shape(Fig 4a), disc shape(Fig 4d), and nodular shape(Fig 4b&c), are observed. The cotton-like shape was observed usually at a low deposition temperature, while nodular morphology was obtained at relatively high temperatures. The disc shape was only observed on the SiC substrate.

The disc shape was only observed on the SiC substrate. When the deposition temperature exceeds 1250°C on the both alumina substrate and SiC substrate, carbon growth becomes unstable. The similar phenomenon was also observed on the effect of the gas pressure. It was found that the deposition process could be easily controlled ifthe C2H2 pressure is lower than 200Torr. The surface quality In addition, of the deposited carbon lines depends on the surface finish of the substrate. depositing carbon on the alumina substrate is easier than depositing on the SiC substrate under the same deposition conditions when the C02 laser was used. These results showed that the processing conditions obviously affect the deposition ofthe carbon.Figure 4 Surface Morphologies ofthe deposited carbon lines using C02 laser on(a) Pelll2 = 200 Torr; Temperature: 800°C; Scanning Speed: O.06mm1see; Alumina substrate; 20 passes.

(b)PolI2 = 100 Torr; Temperature: 12SOoC; Scanning Speed: 0.76mm1sec; Alumina substrate; 20 passes. (c) Poll2 = 100 Torr; Temperature: 1000oC; Scanning Speed: 0.38mm1see; SiC substrate; 20 passes. (d) Pom =200 Torr; Temperature: 900°C; Scanning Speed: 0.13mmlsee; SiC substrate; 20 passes.4853.3 Deposition of insulation layersSince the thermocouple must be electrically insulated from the substrate, an insulating layer is deposited from the gas mixture ofTMS and NH3• It is found that the surface quality of the insulating layer depends on the scan speed and the surface finish of the substrate. For a substrate with poor surface finish, a very low scanning speed is required to get a uniform insulating layer. Under a condition of a 15Torr TMS plus 15Torr NH3 and constant temperature (900°C), a continuous layer with a thickness of about 30Jlm is obtained.

Under a condition of a 15Torr TMS plus 15Torr NH3 and constant temperature (900°C), a continuous layer with a thickness of about 30Jlm is obtained. Electrical measurement showed that it is an insulator.3.4 Electromotive force from SiC/C thermocouplesSome physical parts with the SiC/C thermocouples are shown in Figure 5. The electricalFigure 5 Physical parts with the thermocouples on (a) alumina substrate; (b) the sintered SiC substrate; (c) the single layer SALDVI SiC substrate; (d) multiple-layer SALDVI SiC substrate; (e) multiple-layer SALDVI SiC substrate with the insulating layers at the bottom and top of the thermocouple.486measurements showed that the SALD SiC/C thermocouples, which were deposited on both sintered alumina substrates and SALDVI SiC substrates, exhibited sensitive response to temperature variation.

This value is much lower than Fitterer's results [7]. This difference probably comes from a difference in material structures. First, the SALD carbon line may not be in fully graphite form. In addition, the electrical conductivity of silicon carbide may also be too low. The experiments have shown that SiC obtained is amorphous and non-conductive when it is deposited at low temperature and with an unfocused beam, while the SiC substrates made with the conventional sintering method showed a fully conductive behavior.

These indicated that the low emf constant could be increased through optimization ofSALD process to obtain better structural characteristics of the deposits.It was found thatthe electromotive force constants are almost4 SummaryThe preliminary experiments have shown that it is feasible to make a structure containingin-situ thermocouples by a combined SALDVI and SALD approach.Silicon carbide lines, carbon lines and insulating layers have been deposited from the precursors TMS, acetylene, and a gas mixture of TMS and ammonia respectively. The line profile of silicon carbide changes with temperature and gas pressure. Three type of carbon morphologies, cotton-like shape, disc shape and nodular shape, are identified.It was found that the SALD SiC/C thermocouples respond to temperature variation sensitively. The emf constant of SALD thermocouple is much lower than that obtained with the SiC/C thermocouple made of SiC/C rods.

The emf constant of SALD thermocouple is much lower than that obtained with the SiC/C thermocouple made of SiC/C rods. Optimization of SALD process is needed to improve the emf constant ofthe in-situ SiC/C thermocouple.AcknowledgementThe authors acknowledge the support ofARPA/ONR Grant N00014-96-1-1299 and ONRGrant NOOO14-95-1-0978 in funding this research.Reference: 1. J. L. Maxwell, J. Pegna, D. Messia, and D. DeAngelis, "Direct Feedback Control of Gas-Phase Laser Induced Deposition," SoUd Freeform Fabrication Proceedings, The University of Texas at Austin, 1996, pp227-237 2. O. Lehmann and M. Stuke, "Laser-Driven Movement of Three-Dimensional Microstructures Generated by Laser Rapid Prototyping," Science, Vol. 270, pp1644-1646 3. F.T Wallenberger, "Rapid Prototyping Directly from the Vapor Phase," Science, Vol. 267, pp1274-5 4. G. Zong, R. Carnes, H. G. Wheat and H. L.

270, pp1644-1646 3. F.T Wallenberger, "Rapid Prototyping Directly from the Vapor Phase," Science, Vol. 267, pp1274-5 4. G. Zong, R. Carnes, H. G. Wheat and H. L. Marcus, "Solid Freeform Fabrication by Selective Area Laser Deposition," Solid Freeform Fabrication Proceedings, The University of Texas at Austin, 1990, pp83-90 5. B. R. Birmingham, J. V. Tompkins and H. L. Marcus, "Silicon Carbide Shapes by Selective Area Laser Deposition Vapor Infiltration," SoUd Freeform Fabrication Proceedings, The University of Texas at Austin, 1994, pp348-355 6. 1. E. Crocker, K. 1. Jakubenas, S. Harrion and H. L. Marcus, "SALDVI Optimization for the Tetramethylsilane - Silicon Carbide System," Solid Freeform Fabrication Proceedings, The University of Texas at Austin, 1997, To be published 7. G. R. Fitterer, "A New Thermocouple for the Determination of Temperatures up to at least 1800°C," AIMME Transaction, V105, 1933, pp290-301 8.

G. R. Fitterer, "A New Thermocouple for the Determination of Temperatures up to at least 1800°C," AIMME Transaction, V105, 1933, pp290-301 8. Helena Westberg, Mats Boman, Ann-Sofie Norekrans and Jan-otto Carlsson, "Carbon Growth by Thermal Laser-assisted Chemical Vapor Deposition," Thin SoUd Films, V215, 1992, pp126-133 9. H. O. Pierson, "Handbook of Carbon, Graphite, Diamond and Fullerenes: Properties, Processing and Applications," Noyes Publications, New Jersey, 1993488.

brication of Laser Deposited TiC/Steel Matrix Composite CoatingsW.H. Jiang, R. KovacevicResearch Center for Advanced Manufacturing, School of Engineering, Southern Methodist University, 1500 International Parkway, Suite #100, Richardson, TX 75081, USA Reviewed, accepted August 25, 2003 AbstractThe present work investigates the effect of laser scanning beam speeds and the content of TiC in injected powder on morphologies and microstructures of laser deposited beads of a TiC/H13 tool steel composite. The results show that the beam scanning speeds affect the size and morphology of the beads. During laser processing, TiC melts, decomposes, and subsequently, a number of fine TiC precipitates form during cooling that are uniformly distributed in the tool steel matrix. The beam scanning speeds and the amount of injected TiC exert a strong influence on the morphology and size of the fine TiC precipitates.

The beam scanning speeds and the amount of injected TiC exert a strong influence on the morphology and size of the fine TiC precipitates. It is believed that the precipitated TiC is the primary phase in hypereutectic Fe-TiC. Rapid cooling develops martensite with retained austenite in a steel matrix. The precipitated TiC can refine grains of the steel matrix as a solidified nucleus. TiC/H13 tool steel composite coatings with various contents of TiC were produced using the laser deposition processing technique.IntroductionLaser cladding is widely used for creating various surface coatings of a significant thickness that can effectively protect substrates from harsh service conditions. Various coating materials, such as carbides, oxides, nitrides, borides and their composites, have been developed [1-4].

Various coating materials, such as carbides, oxides, nitrides, borides and their composites, have been developed [1-4]. These coatings display high hardness, wear, erosion and corrosion resistance, heat insulation.Metal matrix composites (MMCs) as coatings have recently attracted much interest, for they combine excellent ductility and toughness of metallic matrices with high strength and hardness of ceramic reinforcements. Due to its high hardness, melting point, and thermodynamic stability as well as availability, TiC is extensively used as a reinforcing phase in MMCs' coatings, such as Ni based [5-8] and steel based composites [8-10]. Due to a relatively low cost, iron and steels as a coating matrix have a potential application prospect. Utilizing laser surface alloying, Ariely et al. [8] produced TiC reinforced steel coatings on the surfaces of Armco iron, AlSl 1045 and 1095 steels, exhibiting a higher hardness. Tassin et al.

Utilizing laser surface alloying, Ariely et al. [8] produced TiC reinforced steel coatings on the surfaces of Armco iron, AlSl 1045 and 1095 steels, exhibiting a higher hardness. Tassin et al. [9] incorporated TiC into the surface of AlSl 316 L stainless steel by laser processing, which substantially improved sliding wear resistance. Jiang and Molian [10] increased life of die-casting dies by laser surface processing with micrometer- and nanometer-sized TiC powder.Slurry erosion happens extensively in the industries of mining, metallurgy, and crude-oil drilling, etc. Materials with improved wear resistance to slurry erosion are in increasing demand in the industries. It is well known that hard materials have a high erosion resistance at a low impact angle, while ductile materials have a high erosion resistance at a high impact angle.524However, in some service conditions, components suffer slurry erosion at varying impact angles.

Therefore, it is of practical significance to develop coatings with an overall excellent erosion resistance, no mater what the impact angles are. Considering their good combination of hardness and ductility, steel matrix composites may be potential erosion resistant materials at various impact angles. Utilizing laser surface processing, the present work tries to develop TiC/H13 tool steel composite coatings as a new erosion resistant material.Laser processing is characterized by its high energy density, low heat input, and consequently, high heating and cooling rates that minimizes its effect on a substrate. However, laser processing parameters such as beam power and beam scanning speed affect metallurgical quality, microstructures, and furthermore, mechanical properties of coatings [10]. In this work, laser deposition is used to synthesize TiC/H13 tool steel coatings with different concentrations of TiC.

In this work, laser deposition is used to synthesize TiC/H13 tool steel coatings with different concentrations of TiC. The effect of beam scanning speeds on the microstructures of coatings is investigated for a given laser beam power and powder feeding rate. Their erosion resistance is being investigated and will be reported.Experimental procedureFig. 1 shows the schematic diagram of the 3D laser cladding system setup used in this investigation. A 1 kW continuous wave Nd: YAG laser with a 200-mm focusing optics is used to deposit the powders. The laser beam guided by the optical fiber is reflected from the partial reflective mirror, and is focused on the substrate by the set of lens arranged in the laser head. The computer-controlled powder feeders are used for dosing and feeding the powder mixture into the desired composition, and allow for an exact setting and continuous change of the powder mixture during the cladding process [11].

The alignment of the laser beam, powder nozzles, and their positioning in relation to the substrate must be accurate for the process to be reproducible. The feeding rate of each of the powders and laser processing parameters are controlled automatically.AISI 4140 steel plate was used as a substrate, whose composition is shown in Table 1. H13 tool steel powder was used as a matrix of composite coatings. The original size of the H13 tool steel powder is around 50-100 µm. Its composition is also shown in Table 1.

1 Schematic diagram of 3D laser cladding system setup.The laser beam power used in the experiments was 380 W, with a spot diameter of 1 mm. The beam scanning speeds were set to 5.08, 7.62, 10.16, and 12.7 mm s-1. A feeding system was used for direct injection of powders with argon flow as a powder carrier. The TiC and H13 tool steel powder was injected by a powder feeder through four nozzles of 1 mm in diameter. The feeding rates of TiC and H13 tool steel are dependent on designed constitutions of coatings. They are (3 × TiC vol.%) g min-1 and (4.65×H13 vol.%) g min-1 for TiC and H13 tool steel, respectively. Single and double layer beads were deposited at various beam scanning speeds. In case of overall coating, substrate samples (25.4 mm × 25.4 mm) were coated uniformly with double layers of about 1 mm in thickness. The beam scanning speed is 10.16 mm s-1.

In case of overall coating, substrate samples (25.4 mm × 25.4 mm) were coated uniformly with double layers of about 1 mm in thickness. The beam scanning speed is 10.16 mm s-1. A 40% overlap of a successive melting track was selected to produce a uniform coating on a substrate.The transverse cross sections of the deposited beads were cut for microstructural examination. The optical microscopy and electron microprobe with energy dispersive analysis of X-ray (EDAX) were used for microstructural observation and microanalysis of composition. X- ray diffraction with Cu Kα was performed to identify the phases in coatings.ResultsLaser beam scanning speeds affect the sizes and morphologies of deposited beads. However, their effect is independent of the constitutions of the deposited powder, that is, the proportion of H13 tool steel to TiC. The typical morphologies of the cross sections of the beads526produced at various beam scanning speeds are shown in Fig. 2.

The typical morphologies of the cross sections of the beads526produced at various beam scanning speeds are shown in Fig. 2. It is well known that a laser clad layer can be divided into three zones, i.e., coating, melting and heat-affected zones in a substrate. From Fig. 2, it can be seen that coating and melting zones are influenced severely by the beam scanning speeds. The faster the beam moves, the smaller the volume and the mass of the coatings. Yet, the melting zone of the substrate decreases with decreasing beam scanning speed, and disappears completely at the minimal speed, i.e. 5.08 mm s-1. It can be seen that only two intermediate beam scanning speeds can develop well geometric and symmetric beads. At the maximum beam scanning speed, 12.7 mm s-1, the coating fails to coincide with the melting zone; while at the minimal speed, the coating tends to lean.

At the maximum beam scanning speed, 12.7 mm s-1, the coating fails to coincide with the melting zone; while at the minimal speed, the coating tends to lean. It is noted that the maximum beam scanning speed leads to involvement of the substrate material into the coating, but fails to homogenize them, as shown in Fig. 2d. This may result from both the larger melting zone in the substrate and higher solidification rate, which mixed much substrate material with deposited materials. Double layer depositions have the same melting and heat affected zones as the single ones. Undoubtedly, a double layer deposition increases the sizes of deposited coatings. Worthy to mention, the beads of the double layer depositions at the minimal beam scanning speed completely lose geometrical symmetry.(a)(b)(c)(d)0.5 mmFig. 2 Morphologies of beads of 60 vol.% TiC/H13 steel formed by single layer scanning at speeds of (a) 5.08, (b) 7.62, (c) 10.16 and (d) 12.7 mm s-1.527From Fig.

2 Morphologies of beads of 60 vol.% TiC/H13 steel formed by single layer scanning at speeds of (a) 5.08, (b) 7.62, (c) 10.16 and (d) 12.7 mm s-1.527From Fig. 2, it can be seen that the interface morphologies between a melting zone and a heat-affected zone change with the beam scanning speeds. By decreasing the beam scanning speeds, their curvatures decrease. In the example of the minimal beam scanning speed, a melting zone disappears and the interface becomes planar. Fig. 3 shows the typical high magnification micrographs of planar (at a speed of 5.08 mm s-1) and curved (at a speed of 12.7 mm s-1) interfaces. Evidently, the curved interface has a better melting bond than the planar.(a)(b)10 µm10 µmFig.

Evidently, the curved interface has a better melting bond than the planar.(a)(b)10 µm10 µmFig. 3 High magnification micrographs of interfaces in 60 vol.% TiC/H13 steel formed by single layer scanning at speeds of (a) 5.08 and (b) 12.7 mm s-1.High magnification observation indicates that TiC particles decompose, and a great number of fine precipitates formed during laser processing. For this reason, the injected coarse TiC particles in beads appear fewer in the low magnification micrographs (Fig. 2). By increasing the beam scanning speeds, the fine precipitates decrease. Fig. 4 shows the typical microstructures of the beads produced at various beam scanning speeds.For all the coatings, the fine precipitates are well distributed in steel matrices. Fig. 5 shows the backscattered electron images of the beads with 40, 60, and 80 vol.% TiC formed at a beam scanning speed of 10.16 mm s-1.

Fig. 5 shows the backscattered electron images of the beads with 40, 60, and 80 vol.% TiC formed at a beam scanning speed of 10.16 mm s-1. In these images, both the coarse TiC and the fine precipitates appear dark, indicating that their constituting elements are lighter than the iron matrix. EDAX demonstrated that fine precipitates are rich in Ti. Furthermore, it can be seen that by increasing the contents of the injected TiC, the fine precipitates increase both in quantity and size. The injected TiC particles exhibit clear, smooth edges and no reaction layer, indicating that they melt and dissolve rather than react with the melt during laser processing.The morphologies and sizes of the fine precipitates are related to the contents of the injected TiC. Fig. 6 shows the fine precipitates in the beads with various contents of injected TiC produced at a beam scanning speed of 10.16 mm/s. In 40 vol.% TiC, they assume substantially spherical and cubic shapes.

In 40 vol.% TiC, they assume substantially spherical and cubic shapes. In 60 vol.% TiC, some precipitates evolve into small flower-petal shapes, while the majority are still spherical. But, in 80 vol.% TiC, the precipitates are rather large and take spherical, cross, rod, and dendrite forms.528(a)Fine precipitates(b)Injected TiC(c)(d)50 µmFig. 4 Microstructures of 60 vol.% TiC/H13 steel formed by single layer scanning at speeds of (a) 5.08, (b) 7.62, (c) 10.16 and (d) 12.7 mm s-1.By a careful examination of Fig. 6b, it can be found that numerous fine precipitates are located at the center of the matrix grains. This indicates that they acted as a heterogeneous site for solidified nucleus of steel matrix grains to nucleate.X-ray diffraction analysis was performed on the coatings with 40, 60, and 80 vol.% TiC produced at the beam scanning speed of 10.16 mm s-1, respectively. The resulting diffraction patterns are shown in Fig. 7.

The resulting diffraction patterns are shown in Fig. 7. It can be seen that the diffraction patterns of all the coatings are similar, and there are only diffraction peaks of steel matrix and TiC, indicating that TiC is only the secondary phase. These results mean that the fine precipitates are also TiC. The steel matrices are martensite and retained austenite.529(a)(b)50 µm50 µm(c)50 µmFig. 5 Backscattered electron images (BEI) of (a) 40, (b) 60 and (c) 80 vol.% TiC/H13 steel formed at a beam scanning speed of 10.16 mm s-1.DiscussionThe present results demonstrate that sizes and morphologies of the deposited beads are closely related to a beam scanning speed. In fact, they are dependent on the proportion of a powder feeding rate to a beam speed for a given powder constituent. During laser processing, laser energy is absorbed by both fed powder and substrate. Undoubtedly, their proportion decides the morphologies of the beads.

During laser processing, laser energy is absorbed by both fed powder and substrate. Undoubtedly, their proportion decides the morphologies of the beads. Changing a beam speed is equal to inversely modifying a powder feeding rate. In the case of a smaller powder feeding rate, or a larger beam scanning speed, the amount of powder injected per unit time is small. Consequently, the energy that the powder absorbs is smaller, and the more energy may be spent to melt a substrate, resulting in a larger melting zone and a smaller bead. That is obvious for the maximum beam scanning speed (12.7 mm s-1) used in the work (Fig. 2d). In the case of a larger powder feeding rate, or a lower beam scanning speed, more powder is injected. Therefore, more energy is absorbed by the injecting powder than the substrate. That results in a larger bead and a smaller melting zone. The situation of the minimal beam scanning speed (5.08 mm s-1) is typical (Fig. 2a).

That results in a larger bead and a smaller melting zone. The situation of the minimal beam scanning speed (5.08 mm s-1) is typical (Fig. 2a). Furthermore, no obvious530melting bond at the interface (at a speed of 5.08 mm s-1) (Fig. 3a) can be explained in this viewpoint.(a)(b)10 µm10 µm(c)10 µmFig. 6 High magnification BEI of (a) 40, (b) 60 and (c) 80 vol.% TiC/H13 steel formed at a beam scanning speed of 10.16 mm s-1.At the minimal beam scanning speed, the lack of a melting zone in a substrate results in a weak interface bond between the coating and substrate, as shown in Fig. 3a. Evidently, this speed is not suitable for producing a coating, not to mention the asymmetrical bead it produced (Fig. 2a and Fig. 3), that may accommodate gas and cause formation of pores during overlapping in coating production. However, the maximum beam scanning speed fails to make the upper coating coincide properly with the lower melting zone.

However, the maximum beam scanning speed fails to make the upper coating coincide properly with the lower melting zone. More severely, some substrate material in the melting zone does not mix well with the deposited materials (Fig. 2d). These disadvantages make it unsuitable for producing a coating. In view of the morphology of beads, both the maximum and minimal beam scanning speeds are not ideal for coating production, while the intermediate beam scanning speeds, 7.62 and 10.16 mm s-1 seem to be appropriate.The smooth edges of the coarse TiC particles indicate that the injected TiC melts and dissolves rather than reacts with the melt. As it is well known, the laser beam is characterized by a high energy density. The degree of absorption of the laser beam on metal and ceramic is very different.

As it is well known, the laser beam is characterized by a high energy density. The degree of absorption of the laser beam on metal and ceramic is very different. Ceramics have a much higher capability to absorb laser energy than metals [12].531Therefore, in spite of its extremely high melting point (3140oC) [13], melting of TiC occurs during processing. As the laser processing time is extremely short, its melting is incomplete and unmelted TiC remains.) 1 1 1 ( C Ti) 0 0 2 ( C Ti) 0 1 1 (e t i s n e t r a M) 1 1 1 (e F - γ) 0 2 2 ( C Ti) 0 0 2 (e t i s n e t r a M) 1 1 3 ( C Ti) 0 2 2 (e F - γ) 1 1 2 (e t i s n e t r a M) 2 2 2 (e F - γ) 0 2 2 (e t i s n e t r a M80 vol.% TiCy t i s n e t n I60 vol.%TiC) 0 0 2 (40 vol.%TiCe F - γ304050607080901002θ, degFig.

7 X-ray diffractogram of 40, 60, 80 vol.%TiC/H13 tool steel formed at a beam scanning speed of 10.16 mm s-1.As melted TiC has a high tendency to dissociate into Ti and C, melting of a large amount of TiC results in a substantial increase in the concentration of Ti and C in the melt. Undoubtedly, TiC reprecipitate in the melt during cooling. EDAX and X-ray diffraction analysis confirms that fine TiC reprecipitated during subsequent cooling may contain less amount of other carbide forming elements, such as Mo and Cr. As the melt cools at a relatively high rate, solidified structures must be in a thermodynamic nonequilibium. That results in the development of various morphologies of TiC. At a higher beam speed and a lower content of injected TiC, the TiC precipitated in globular and cubic shapes. By lowering the beam scanning speed and increasing the amount of injected TiC, they tend to assume hexagonal and flower-petal shapes and finally the form of dendrites.

By lowering the beam scanning speed and increasing the amount of injected TiC, they tend to assume hexagonal and flower-petal shapes and finally the form of dendrites. This shows that the morphologies of TiC are closely related to the processing parameter and composition of melts. Considering the high contents of injected TiC, other alloying elements in the melt are neglected. The melt is approximately taken as the binary components of Fe and TiC. Refer to the binary Fe-TiC phase diagram (Fig. 8) [14], we explain qualitatively the morphologies of TiC. It can be seen that all the present constitutions of the coatings are hypereutectic. Therefore, TiC precipitates as a primary phase during cooling. Their morphologies depend on growth, as their nuclei are globular. Various morphologies of TiC can be developed during growth [15]. At a lower beam speed, solidification lasts a longer time and532primary TiC has time to grow a bit.

Various morphologies of TiC can be developed during growth [15]. At a lower beam speed, solidification lasts a longer time and532primary TiC has time to grow a bit. For a higher content of injected TiC, the melt has a higher concentration of Ti and C, that is, a higher supersaturation. Therefore, their growth rate is higher, as a high solute concentration speeds up growth. Observed morphologies of TiC, spherical, faceted, rod, flower-petal shapes, and dendrites are from various growth stages. Due to rapid solidification, it is believed that the deposited beads are far from equilibrium and precipitated TiC is only primary phase during solidification.170016501600T, oC1550L1500Fe+L145014001380806040 vol.%TiC1350TiC+Fe13000 TiC10 20 30 40 50 60 70 80 90 100Feat.% FeFig. 8 Fe-TiC phase diagram [14].In all the coatings, fine TiC precipitates are uniformly distributed in the matrix. They do not segregate around coarse TiC particles.

8 Fe-TiC phase diagram [14].In all the coatings, fine TiC precipitates are uniformly distributed in the matrix. They do not segregate around coarse TiC particles. This indicates that titanium and carbon atoms released by injected TiC are well distributed in melts. Their homogeneous distribution must be attributed to vigorous convection caused by rapid heating of the laser beam. A convective trace is evidenced in Fig. 2. It seems that the larger the beam scanning speed, the more obvious the convection.Rapid cooling leads to the transformation of steel matrices into martensite (Fig. 7) that hardens a steel matrix. The fine precipitated TiC was observed to be a solidified nucleus for the steel matrix (Fig. 6b). They refined the grain structures of the matrix. In fact, it is well known that TiC is a good refiner for steels.

6b). They refined the grain structures of the matrix. In fact, it is well known that TiC is a good refiner for steels. Undoubtedly, this favors to modify solidified structures and furthermore, to improve both the strength and toughness of the steel matrix.The present work indicates that microstructures of deposited beads are closely related to processing parameters. The beads with microstructures of coarse injected and fine precipitated TiC are developed. It is believed that tailored microstructures of TiC/H13 steel composite coatings can be achieved by choosing appropriate laser processing parameters.533As well known, fine particles would improve the mechanical properties of composites. It is expected that melting of coarse TiC and subsequent reprecipitation of fine TiC would impart better overall mechanical properties to coatings than the mechanical insertion of coarse TiC particles.

Both the coarse remaining and the fine precipitated TiC particles are believed to have a clean interface with the steel matrix, as the original surface of injected TiC melts, a fresh surface is developed, and fine TiC precipitates are formed in-situ within the melts. Therefore, a stronger interface bond between TiC and the matrix may be achieved. As well known, the interface is a crucial location affecting the mechanical properties of composites. Impurity atoms usually tend to adhere to the surfaces of particles that disfavor an interface bond. In-situ formed composites can circumvent such disadvantages and therefore, have received much attention recently [16]. It is believed that the interface quality between TiC and the matrix in the present coatings is the same as that in in-situ formed composites. It is expected that the coating composites possess better overall mechanical properties.Conclusions1. The beam scanning speed affects the size and morphology of the deposited beads.

It is expected that the coating composites possess better overall mechanical properties.Conclusions1. The beam scanning speed affects the size and morphology of the deposited beads. The faster the beam moves, the larger the melting zone at the substrate, and the smaller the deposited zone. But, the size of heat-affected zones is independent on beam scanning speed.2. The higher beam scanning speed fails to mix the deposited material properly with a substrate. 3. During laser processing, TiC melts, decomposes, and subsequently, a number of fine TiC precipitates form. They are uniformly distributed in the steel matrix, which is attributed to vigorous convection caused by rapid heating.4. The beam scanning speeds and the amount of injected TiC have a remarkable impact on the morphology and size of the precipitated TiC. The lower the beam scanning speeds and the more the injected TiC, the larger the precipitated TiC.

The lower the beam scanning speeds and the more the injected TiC, the larger the precipitated TiC. The larger TiC assumes spherical, rod, flower-petal, and dendrite shapes, while the smaller TiC is spherical and faceted.5. It is believed that precipitated TiC is the primary phase in hypereutectic Fe-TiC. 6. Steel matrices in coatings are martensite with retained austenite. Precipitated TiC can refine grains of the steel matrix as a solidified nucleus.AcknowledgementsThe authors would like to express gratitude to Dr. H. H. Mei, Mr. Z.G. Liu and Mr. M.Valant for their helps during the experimental work.References1. T.C. Lei, J.H. Ouyang, Y.T. Pei and Y. Zhou, Surface Engineering, 12, 1996, 55. 2. B.J. Kooi, Y.T. Pei and J.Th.M. De Hosson, Acta Mater., 51, 2003, 831. 3. A. Agardwal, Surface Engineering, 17, 2001, 66. 4. Y.T. Pei, J.H. Ouyang and T.C. Lei, Surface and Coating Technology, 81, 1996, 131. 5. Y. T. Pei and T.C. Zou, Mater. Sci. Eng. A, A241, 1998, 259. 6. Q.

viewed PaperFabrication of Liquid-Filled Voronoi Foams for Impact Absorption Using Material Jetting Technology Tosh Kāneala Brown-Moore†, Srivatsan Balaji‡, Tyler Williams‡, Jeffrey Lipton‡*† Department of Materials Science and Engineering, University of Washington, Seattle, WA‡ Department of Mechanical Engineering, University of Washington, Seattle, WACorresponding author. Email: jilipton@uw.eduAbstractAn important consideration in the design of any mechanical system is its ability to isolate and dissipate vibrational and impact energy. Closed-cell foams utilize cell crumpling to absorb energy, relying solely on viscoelastic effects for damping. Liquids, however, can generate large amounts of damping from fluid channel friction and turbulence.

Liquids, however, can generate large amounts of damping from fluid channel friction and turbulence. We produced closed-cell foams that are liquid filled, resulting in tunable materials that absorb energy better than either component on their own, using a Voronoi generation model and a J750 printer that could jet curable and incurable liquids. We found that by changing the wall thickness and liquid percentage, we achieve a stiffness range of 4.1 N/mm to 80 N/mm. Our work introduces this new class of damping metamaterial that can absorb tunable amounts of energy per unit volume. These impact-absorbing structures may benefit applications such as protective equipment, healthcare, and automotive industries.IntroductionImbalances inherent to turning components in mechanical systems are often the origin of vibrations that can excite natural frequencies leading to long term performance reduction, reduced service life, and human discomfort.

[1,2] Passive vibration mitigation employs dampers and isolators to attenuate a vibration source, preventing its propagation to the rest of the system and therefore protecting it from damages. Passive dampers are used to dissipate incoming kinetic energy and convert it to heat, usually involving viscous fluids, viscoelastic materials, piezoelectric elements, or electromagnetic devices. [3,4,5] Conversely, passive isolators such as closed-cell foam slabs, metal coil or wave springs, wire rope isolators, and rubber machine mounts are used to decouple a vibration source from a mechanical system, separating a system’s natural frequency from the excitation frequency therefore avoiding resonance. [6,7]Nonperiodic transient shock loads also commonly initiate vibration. Isolators store shock energy and gradually dissipate it via free vibration, requiring large displacements to be efficient. Dampers utilize friction to dissipate energy but limit displacement, reducing isolation efficiency.

Dampers utilize friction to dissipate energy but limit displacement, reducing isolation efficiency. These competing mechanisms make it difficult to design a system where isolators and dampers work in tandem. [8] Furthermore, traditional elastomeric isolators can only operate within a predetermined narrow band width of harmonics. [9] Fluid viscous dampers can combat harmonics653in a larger range of frequencies but are anisotropic and lack compliance. Additive manufacturing enables the creation of metamaterials capable of both damping and isolation on a tunable spectrum.Figure 1. Visualization of vibration attenuation mechanismsUsing the Stratasys J750, we created a liquid filled closed-cell foam that can dissipate energy through fluid channel friction and isolate vibration via cell crumpling.

This new metamaterial mitigates vibration while maintaining compliance and geometric freedom, utilizing the interface between the two phases and the inherent damping properties of liquid to achieve a higher attenuation density than traditional dampers or isolators on their own.Figure 2. Schematic representation of metamaterial constructionTo generate these Voronoi foams, a Voronoi-based infill-generating algorithm was used [10]. This algorithm functions by generating the seed points to fill the volume of a given 3D model, using the Parallel Jump Flood algorithm to indicate the volumes of the Voronoi cells, then running through several post-processing steps to find the cell walls and apply the desired wall thickness. [10,11] The properties of foams generated in this manner depend on seed-point distribution, seed-654point density, shell thickness, shell shape, overall size, wall thickness, and liquid percentage relative to total volume.

Seed point distribution and density were determined probabilistically, with each voxel in our discretized space as a probability of 5% of generating a seed point at its location. Shell thickness was kept at a constant 1.83mm, our tested minimum to prevent leakage during printing. With a rectangular prism shell shape, shell size was kept a constant 40 mm by 40 mm by 10 mm. Wall thickness was altered and tested in a range of 0.138mm to 0.975mm, including a baseline solid sample test with a technically infinite wall thickness. Likewise, liquid percentage relative to total volume was altered and tested in a range of 37% to 53%. Further studies may alter parameters we kept constant to obtain and explore a broader range of material properties.It would be impossible to make a closed-cell liquid-filled foam with traditional manufacturing methods.

The design freedom granted by additive manufacturing enables the creation of such a material that maintains the strengths of both traditional dampers and isolators in a tunable manner.ResultsLimit of ParametersFigure 3. Effect of wall thickness on print successThe Stratasys J750 has a 7° vertical angle tolerance. This constraint can be somewhat bypassed when liquid is used as the pseudo-support material, but only up to a maximum cavity width of 10mm. [12] Print success is also dependent on the thickness of cell walls between seed points. Walls of insufficient thickness are not strong enough to hold the weight of liquid in higher layers, which can result in delamination or total failure. When layer delamination occurs, the bulk solid portion will partially or even completely remain intact but liquid leaks out of the foam depending on the severity of delamination.

When layer delamination occurs, the bulk solid portion will partially or even completely remain intact but liquid leaks out of the foam depending on the severity of delamination. While not explored here, leakage could be purposefully employed to successfully print near closed-celled foams with air cavities that would be normally impossible with the 7° angle constraint.655Quasi Static CompressionFigure 5. Load behavior for foam of varying wall thicknessesRegardless of the perceived success of a print, all foam configurations were subject to a quasi-static compression test to gauge their fundamental failure behavior. As each foam was compressed to failure, liquid gradually leaked out of the sides. Lower wall thickness foams have a smoother trend at higher displacements because its behavior is dominated by the larger amounts of incompressible liquid present. The thinner walls are easily broken to allow liquid to flow.

The thinner walls are easily broken to allow liquid to flow. Higher wall thickness foams are less stiff at higher displacements because the dominating mechanism is compressive densification of the elastomer. Errors become more pronounced at higher wall thicknesses because it is more difficult for liquid to leak out of the more robust cells; once the forces from the liquid exceed the capacity of the thicker walls the leakage is much more violent causing more perturbations in the data.VibrationFigure 6. Foam behavior as a function of frequency656The stiffness of each foam configuration also changes as a function of excitation frequency. Lower wall thickness foams have a more gradual increase in stiffness, maintaining it over a broader range of frequencies, while larger wall thickness foams including the solid baseline have a very sudden increase in stiffness at higher frequencies.

This supports the idea that viscous dampers operate in a wider frequency range than traditional elastomeric dampers. [13] Furthermore, if there is too much solid material, stiffness rises which inhibits the foams’ ability to displace therefore decreasing its isolation efficiency. Cell crumpling and liquid friction surface area are both limited at higher thicknesses but don’t leak as easily. This tradeoff would need to be optimized depending on the application and required specifications.Figure 7. How dissipation potential changes with vibration frequencyIn purely elastic materials, stress and strain are in phase. In purely viscous materials stress and strain are offset by a phase lag of δ = 90°. By this logic, as the volume of liquid within the foam increases and the wall thickness decreases, phase lag should approach δ = 90° and tan(δ) should tend upward toward infinity.

By this logic, as the volume of liquid within the foam increases and the wall thickness decreases, phase lag should approach δ = 90° and tan(δ) should tend upward toward infinity. As wall thickness increases, which in turn decreases the volume of fluid within the foam, phase lag should approach δ = 0 and tan(δ) should likewise approach zero. Foams of higher wall thickness should have a lower tan(δ) because they have a higher density of elastic material. However, this is not always the case. At any given frequency, the phase lag order is not consistent with what would be intuitively expected. We attribute this behavior to the fact that liquid cells of different sizes resonate at different frequencies. Tan(δ) is the ratio of the loss modulus to the storage modulus and provides an indication of how well a material can dissipate energy. A higher tan(δ), inherent to more viscous materials, suggests that a material is more well suited for energy dissipation rather than storage.

A higher tan(δ), inherent to more viscous materials, suggests that a material is more well suited for energy dissipation rather than storage. Liquid cells of different sizes have different resonant frequencies, explaining the seemingly unintuitive ordering.657Figure 8. Alternative independent variable to wall thicknessThe ability of a foam to efficiently dissipate energy, indicated by high tan(δ), generally increased as a function of liquid percentage. This is defined as the ratio of liquid volume to the total volume of foam. At higher frequencies such as = 22.93 rad/s this isn’t the case, attributed once again to the fact that different foam configurations have different resonant frequencies. The foam with a liquid percentage between 35-40% had a natural frequency that destructively interfered with the excitation source thus dissipating energy more efficiently.

The foam with a liquid percentage between 35-40% had a natural frequency that destructively interfered with the excitation source thus dissipating energy more efficiently. [14]⍵ConclusionAn important consideration in the design of any mechanical system is its ability to isolate and dissipate vibrational and impact energy. Additive manufacturing enables the creation of metamaterials that can isolate and damp vibration and shock in tunable amounts. Using a Voronoi generation model and a Stratasys J750 inkjet 3D printer, we created liquid filled closed-cell elastomeric foams with a stiffness range of 4.1 N/mm to 80 N/mm depending on the wall thickness of the printed specimen. This new class of damping metamaterials utilizes the interface between solid and liquid phases to dissipate energy and traditional cell crumpling as an isolation mechanism.

This new class of damping metamaterials utilizes the interface between solid and liquid phases to dissipate energy and traditional cell crumpling as an isolation mechanism. Energy is attenuated while maintaining compliance and complete geometric freedom, opening the door to applications in the protective equipment, healthcare, and automotive industries. There are still many aspects of this new material that can be explored including seed point position, liquid channels, and varied liquid percentage that would further improve its damping and isolation capabilities.658Bibliography1. H. Mahdisoozani et al., “Performance Enhancement of Internal Combustion Engines through Vibration Control: State of the Art and Challenges,” Applied Sciences, vol. 9, no. 3, Art. no. 3, Jan. 2019, doi: 10.3390/app9030406.2. P. Jönsson and Ö. Johansson, “Prediction of vehicle discomfort from transient vibrations,” Journal of Sound and Vibration, vol. 282, no. 3, pp. 1043–1064, Apr.

It can be used for the fabrication of new components but also repair applications. So far, volume build-ups were mostly carried out with one single material only. However, loading conditions may strongly vary and, hence, the use of more than one material in a component would yield major benefits. By means of multi-material build-ups, cost-intensive alloys could be used in highly-loaded areas of the part, whereas the remaining part could be fabricated with cheaper compositions. The selection of combined materials strongly depends on the requested thermo-physical and mechanical properties. Within this contribution, possibilities of material combinations by LMD and selected examples of beneficial multi-material use are presented.IntroductionIn several industrial branches, Laser Metal Deposition (LMD) is already a common technology for coating applications, part refurbishment but also for manufacturing of functional components.

Due the process’ suitability to a broad spectrum of materials (e.g. Fe-, Ni-, Co- or Ti-based alloys) LMD is used for a large range of products such as components for jet engines, turbines, tooling or medical implants. The LMD is a nozzle based manufacturing process (see Figure 1). The powdery material is blown into the process zone by the nozzle. The laser-powder interaction causes a preheating of the powder particles. The preheated particles are then finally absorbed by the induced melt pool and deposited on the substrate.mixing ratio, an in-situ alloying process without any additional joining can be achieved by LMD [1…5].Material combinations such as Ti-Ta, Fe-Ta or Fe-Cu are highly interesting for industrial applications but cause several issues due to thermo-physical properties, poor differing miscibility or differing absorptivity.

Therefore in depth understanding of process an stress conditions gradients), material properties (micro- and macrostructure) failure mechanisms is obligatory for an industrial transfer. In order to counter these problems the following approaches for layer wise multi- material build-up by LMD with powder shall be investigated within this paper:(temperaturefields,andoccurringBefore being blown into the process zone by the nozzle, different powder materials can be mixed in-situ by an integrated powder-mixing chamber (see Fig. 1).

1). Due to the possibility of a continuous as well as localized control of theI.II.sharp material transition of binary material combinations, graded material transition of binary material combinations.2530Figure 1: Powder-based multi-material LMD processing [6]I Sharp material transition of binary material combinationsTi6Al4V – Ta: Ti6Al4V exhibits good mechanical as biocompatibility and is therefore a commonly used material for the fabrication of implants. Due to better osseiointegration properties and biocompatibility tantalum can be used for improving the integration of implants into the human body. Therefore combining tantalum and Ti6Al4V could lightweight implants with good mechanical properties and an increased lifetime [7].propertiesas wellleadtoNevertheless, joining Ti6Al4V and tantalum raises the issue of a large gap between the melting ranges of the materials (see Table 1).

Due to the high melting temperature of tantalum the simultaneous processing of these materials may cause local evaporation phenomena, which may lead to the formation of pores. By using an active2531the Ti6Al4V substrate can be cooling prevented from evaporation whilst fully melting tantalum powder material. Furthermore since Ti6Al4V as well as tantalum are susceptible to oxidation at high temperatures processing has to be conducted using a shielding gas atmosphere.theTable 1: Mechanical and Thermo-physical properties of Ta and Ti6Al4V [8, 9]Density [g/cm³] Yield Strength [MPa] CTE at 20 °C [10E-6/K] Melting Range [°C] Evaporation Temperature [°C]Ti6Al4V 4,45 828 8,9 1604-16603260Ta 16,6 170 6,5 29965458As can be seen in the binary phase diagram (see Figure 2) titanium and tantalum show good miscibility at low temperatures and there is no formation of intermetallic phases.

Furthermore due to similar CTE the bonding zone is less susceptible to thermally induced cracking.Figure 2: Ti-Ta binary phase diagram [10]stated processing and tantalum were successfully joined by LMD. A transition from Ti6Al4V to tantalum without any cracks and almost no pores could be achieved (see Figure 3). The high processingTakinginto accountthecharacteristics Ti6Al4Vtemperatures induced a large melt pool.

The high processingTakinginto accountthecharacteristics Ti6Al4Vtemperatures induced a large melt pool. Due to the the melting temperatures and convection phenomena within the melt pool partial solidification of high melting tantalum areas within a lower melting matrix can be observed (see Figure 3 and Figure 4)large gap betweenTantalumAm Transition Zone400 µmmTi6Al4 VFigure 3: Metallographic image of the LMD transition zone from Ti6Al4V to TaAmB300 µmFigure 4: SEM Image of the LMD transition from Ti6Al4V to Ta; BSE analysis of marked regionA BSE analysis of the transition zone showed the formation of areas with a broad range of differing chemical compositions (see Figure 4 and Figure 5). A significant characteristic of the transition zone is the formation of two major phases. The tantalum content of first phase (see Figure 4 and Figure 5, spectrum 1, 5) is above 80 at. %. Additionally ßTi, Al, and V can be found within this phase.

The tantalum content of first phase (see Figure 4 and Figure 5, spectrum 1, 5) is above 80 at. %. Additionally ßTi, Al, and V can be found within this phase. The microstructure of this2532phase is homogenous, which could be a result of the short solidification time due to a small melting range. The second phase (see Figure 4 and Figure 5, spectrum 7, 11, 15) shows a tantalum content of less than 80 at. %. Moreover, this phase also consists of αTi, Al, V and exhibits a very fine dendritic microstructure. This microstructure might be caused by the increased solidification time due to a larger melting range. Furthermore, due to the high melting temperature (3020 °C) of pure tantalum small (diameter approx. 5 µm) globular pure tantalum areas we formed within a matrix with a higher content of Ti, Al and V (see Figure 6 and Figure 7, spectrum 19, 20).10080%60t h g i eW4020015 No.

5 µm) globular pure tantalum areas we formed within a matrix with a higher content of Ti, Al and V (see Figure 6 and Figure 7, spectrum 19, 20).10080%60t h g i eW4020015 No. of Spectrum Ti V Ta Al71115Figure 5: Results of BSE analysis of the investigated areas 1-15B100 µmFigure 6: SEM Image of LMD Ti6Al4V/tantalum transition; BSE analysis of marked region%t h g i eW100 90 80 70 60 50 40 30 20 10 01920212223No. of SpectrumAlTi V TaFigure 7: Results of BSE analysis of the investigated areas 16-23Tantalum - SS AISI 316L: Besides the use for medical application due to the high melting temperature and excellent corrosion resistance tantalum is also highly attractive for applications in the chemical and energy industry.

Since the procurement of tantalum is three dimensional multi- very expensive material in combination with low cost corrosion resistant materials such as stainless steel could enable new cost efficient design and manufacturing possibilities.processingoftantalumIn this paper, the deposition of SS 316L on commercially pure tantalum shall be presented. The multi material processing of tantalum and SS 316L raises three major issues. As stated above due to the high melting temperature of tantalum the simultaneous processing of tantalum and lower melting local evaporation materials may cause phenomena, which may lead to the formation of pores. Secondly tantalum and SS 316L show a significant difference in CTE. This may lead to thermally induced crack within the bonding zone. (see Table 2)Furthermore, as can be seen in the binary Fe-Ta phase diagram in Figure 8, iron and tantalum only have a limited miscibility and tend to the formation of intermetallic2533phases.

(see Table 2)Furthermore, as can be seen in the binary Fe-Ta phase diagram in Figure 8, iron and tantalum only have a limited miscibility and tend to the formation of intermetallic2533phases. Especially the Laves structure ε phase (Fe2Ta) may cause an embrittlement of the bonding zone.Table 2: Mechanical and Thermo-physical properties of Ta and SS 316L [10, 11]Density [g/cm³] UTS [MPa] Yield Strength [MPa] CTE at 20 °C [10E-6/K] Melting Range [°C]SS 316L 7,99 558 290 15,9 1304-1371Ta 16,6 285 170 6,5 2996Figure 8: Binary Fe-Ta phase diagram [12]400 µmFigure 9: Metallographic image of SS 316L deposited by LMD on tantalum substrateFor ensuring a full melting of the tantalum substrate the LMD process wascombined with simultaneous conduction heating of the substrate.

The process parameters were optimized in order to ensure low evaporation of the SS 316L whilst fully melting the tantalum and in order to prevent the from oxidation at high tantalum temperatures.As can be seen in Figure 9, a dense transition from tantalum to SS316 L and a high fusion could be achieved. degree of this Nevertheless, as mentioned before, material combination is very susceptible to thermally induced cracking due to differing thermal expansion. The highlighted areas in Figure 9 show the formation of micro cracks along the transition zone. Beside the large CTE difference the formation of intermetallics may be another reason for the formation of cracks. very 10 inhomogeneous material distribution within the the presence of different metallic phases.

very 10 inhomogeneous material distribution within the the presence of different metallic phases. In order to investigate the single phases BSE analysis shall be conducted.Figureshowsatransition zone, whichindicates100 µmFigure 10: Metallographic image of SS 316L deposited by LMD on tantalum substrateAISI SS3 04L – Copper: Due to its excellent thermal and electrical conduction but poor mechanical properties, integrating copper into multi-material systems is highly relevant for industrial applications. Therefore joining stainless steel, which provides beneficial2534mechanical properties, and cooper shows high potential for industrial applications such as conductor tracks, heat exchangers or coatings on thermal-loaded parts [13].Figure 11: Wavelength dependence of the absorption rate of different metalsJoining copper to stainless steel causes two major issues: Due to the great difference in NIR absorptivity between copper (approx.

2 %) and stainless steel (approx. 50 %) a high level of power is required to overcome the low absorptivity of copper and to fully melt both materials (see Figure 11). Furthermore due to the higher temperatures the sudden temperature jump while melting may lead to an overshooting and blow out of molten material. Therefore, a green laser source with a wavelength of 532 nm was used.increasingabsorptivityatSecondly in Figure 12 can be seen that the binary copper-iron system shows limited miscibility with two peritectic points. One peritectic point is on the copper rich side at 1096 °C with 3.5 % miscibility of iron in copper. The second one is on the iron rich side at 1485 °C with a maximum miscibility of 7.5 % in γ-Fe.Figure 12: Fe-Cu binary phase diagram [14]parameters were optimized in order to obtain non-porous volumes and low penetration of the filler material (see Figure 13).

The microstructure of characterized by the is inhomogeneous distributed globules in a copper-rich matrix (see Figure 14 and Figure 15). By EDX analysis the composition of the copper track was determined. The copper-rich matrix exhibits a composition with 3.5 at. % Fe. This approves with the binary phase diagram in Figure 12. Once the melt pool temperature solidus temperature first regions with a higher amount of Fe start to solidify and form globular areas due to surface tension. Therefore the amount of Fe in the melt decreases until the peritectic point at 96.5 at. % Cu is reached. At this point the remaining melt solidifies and forms the matrix with peritectic composition. Due to decreasing miscibility of iron in copper with sinking temperature iron rich segregations are likely to occur at the grain borders.

Due to decreasing miscibility of iron in copper with sinking temperature iron rich segregations are likely to occur at the grain borders. The uneven distribution of the globules within the matrix the Marangoni- can be explained by convection, which is driven by thermally and/or concentration induced surface tension gradients. The chemical composition of the globules is close to the iron-rich peritectic composition with 11 at.

The chemical composition of the globules is close to the iron-rich peritectic composition with 11 at. % Cu.Theoperatingsingletrackfallsbelowthe2535500 µmFigure 13: LMD copper track on SS 304 L50 µm20 µmFigure 14: SEM of the copper track]%[n o i t c a r fs s a m100 90 80 70 60 50 40 30 20 10 0CrFeNiCuMeasurement AMeasurement BFigure 15: EDX spectrum of zones A and B of FigureII Graded material transition of binary material combinationsSS AISI 316L – INCONEL 718: As already mentioned the joining of multi- materials may cause severe stresses within the transition zone due to differing thermos- physical and mechanical properties. Therefore a graded transition (see Figure 12) within thebonding zone is a promising approach for decreasing stress gradients and reducing cracking.

Therefore a graded transition (see Figure 12) within thebonding zone is a promising approach for decreasing stress gradients and reducing cracking. Within the graded transition from SS AISI 316L to the nickel- based superalloy INCONEL 718 shall be presented (see Figure 16 and Figure 17).this workFigure 16: Functional principle of graded LMD transitions2 mm µmFigure 17: Metallographic image and EDX analysis (Fe, Ni, Nb) of the generated AISI 316L-INC718 graded material transitionBased on two layers of SS 316, deposited on a SS 304L substrate, the composition of the powder mixture used for2536building up the transition was linearly varied in order to gradually replace the amount of SS 316L by INCONEL 718. The metallographic image in Figure 17 shows that a dense and almost crack free volume with only a little amount of pores was achieved. By EDX analysis the chemical transition process was investigated (see Figure 17). Since there is no Nb in SS 316L but approx.

By EDX analysis the chemical transition process was investigated (see Figure 17). Since there is no Nb in SS 316L but approx. 5.0 weight % in INCONEL 718 Nb was used as a marker element for detecting the gradual transition.The EDX analysis, which was conducted at the layers 1, 6, 10, 14 and 19 shows that in the bottom layer a crack free deposition of SS 316L was achieved. Starting in layer 3 the feeding rates were adapted linearly in order to accomplish a linear transition from SS 316L to INCONEL 718. The linear growth of the Nb mass fraction, which was measured by EDX, indicates a linear transition from SS 316L to INCONEL 718 (see Tab. 3).

The linear growth of the Nb mass fraction, which was measured by EDX, indicates a linear transition from SS 316L to INCONEL 718 (see Tab. 3). Furthermore, the EDX analysis of layer 19 confirms that the SS 316L has been fully replaced by Inconel 718.Table 3: Chemical composition of the generated AISI 316L-Inconel 718 graded material transition (EDX)LayerFe [%]Ni [%]Cr [%]Mo [%]Mn [%]Inconel 718rest50.0- 55.017.0- 21.02.8- 3.30.351918.450.619.73.801430.640.719.53.30.41038.732.919.33.60.7656.119.218.83.11.1165.511.018.92.61.4SS 316Lrest10.0- 13.017.5- 19.52.0- 2.50.0- 2.0Nevertheless, the layer wise modification of the powder mixture composition also raises the issue of ensuring a homogenous materialNb [%]4.7 - 5.564.43.61.20.00.0distribution within each layer.

An uneven material mass fraction distribution may cause and inhomogeneous microstructure an therefore differing and thermo-physical mechanical properties within a single layer.The metallographic in Figure 18 confirms that within a single layer different micro structures can be found. Figure 18 C shows a coarse columnar grain structure whereas Figure 18 A exhibits a fine and rather equiaxed grain structures. This indicates the in situ alloying of different metallic or intermetallic phases in dependence of the local mass distribution. Moreover, according to Figure 18 B local micro cracking along the grain orientation has occurred. This may indicate a local embrittlement due to undesired local phase formation.imagefractionBCA200 µmFig.

This may indicate a local embrittlement due to undesired local phase formation.imagefractionBCA200 µmFig. 18: Metallographic image of the generated SS 316L-INCONEL 718 material transition (layer 6)Conclusion and OutlookWithin this paper, the three dimensional processing of multi-materials by Laser Metal Deposition, related benefits and resulting major issues are presented. Therefore this work gives an overview for LMD multi- material combinations with high industrial relevance. following materials The combinations were investigated:Ti6Al4V – Tantalum2537SS 316L – Tantalum - SS 304L – CP Copper - SS 316L – INCONEL718Moreover approaches were applied:thefollowingprocessingSharp material transition of binary material combinationsGraded material transition of binary material combinationsThe investigations on LMD processing Ti6Al4V have shown that a dense, crack free material transition can be achieved.

Due to the high melting temperature of tantalum a high degree of fusion could be observed. Furthermore the good miscibility at low temperatures and the high degree of fusion causes the formation of locally varying phases with differing microstructures.The multi-material processing of SS 316L and tantalum clearly showed that by using substrate heating a dense transition with a sufficient be of accomplished. Nevertheless the major issues of differing CTE formation of intermetallics caused micro cracking within the transition zone. Therefore different ways for avoiding these phenomena shall be investigated within further research.degreefusioncanandThe successful multi-material processing of the material combinations Ti6Al4V/Ta and SS 316L/Ta might be a promising approach for the dissimilar joining of Ti6Al4V and SS 316L using tantalum as interlayer material.

This approach shall be investigated within further research.Moreover an analysis of a SS 304L and CP copper transition zone manufactured by LMD was conducted. Due to poor miscibility and a high melting range the transition zone shows a composite structure consisting of iron rich globules within a cooper rich matrix. Due to the peritectic behaviour of the Fe-Cu materialsystem the matrix exhibits a Fe atomic fraction of 3.5 at. %. The globules characteristic distribution within the peritectic matrix may result from the occurring Marangoni effect.The great number of different material combinations and, moreover, the broad range of industrial applications shows the high potential of the LMD process for multi- material fabrication.

brication of Multi-Layered Carbon Structures Using LCVDRyan W. Johnson, Scott N. Bondi, Daniel L. Jean, Chad E. Duty, MingXuan Jiang, and W. Jack LackeyRapid Prototyping and Manufacturing Institute Woodruff School of Mechanical Engineering Georgia Institute of Technology Atlanta, GA 30332AbstractOthers have used Laser Chemical Vapor Deposition (LCVD) to create 3-D fibrous structures and helical springs. Current research efforts focus on the creation of more advanced three-dimensional carbon objects through the use of multi-layered deposition. Multi-layered structures require an understanding of interlayer adhesion and the propagation of geometric anomalies through multiple layers. An important aspect in minimizing these shape anomalies is the implementation of closed loop temperature control. Several laminated carbon structures are presented with discussions and observations about the fabrication process and visual characteristics of each.

Several laminated carbon structures are presented with discussions and observations about the fabrication process and visual characteristics of each. The major issues in using LCVD to create multi-layer carbon structures are addressed.IntroductionNumerous previous attempts at constructing layered deposits using LCVD have been made. Such efforts have typically produced non-uniform periodic deposits or dimensionally unstable structures.1 There is a belief that the general nature of the LCVD process is responsible for these irregularities. The existence of substrate protrusions and the creation of abnormalities in one layer of a deposit are thought to be amplified in each successive layer. This behavior stems from localized heating of surface irregularities causing increased local deposition rates.

This behavior stems from localized heating of surface irregularities causing increased local deposition rates. Online monitoring and control of the reaction is necessary to enable consistent and uniform deposition.Recent efforts have been made to determine the feasibility of building multi-layered pyrolytic carbon structures using a temperature controlled LCVD process. All structures built for this objective were created using a multiple pass process whereby the substrate was either translated along one axis in alternating directions of travel or in a continuous circular pattern. The resulting structures were multi-layered carbon walls or cylinders conceptually similar to those presented by Messia, et al.2 Theoretically, this construction technique will create layers that are intimately bonded producing a final structure that could exhibit either bulk solid or laminated solid properties.

Many observations were made during the construction of the walls with regard to a new temperature control system and the general nature of the deposits. Some deposits were cut and polished so as to provide a view of the cross section of the walls and thus allow inspection of their integrity. Analysis of the cross sections and other features of the walls provide insight into the use of LCVD to create multi-layered structures.61Experimental SetupGeorgia Tech’s LCVD reactor is powered by a 100 W CO2 laser with a 200 µm diameter spot size. The system incorporates a fixed laser and reagent supply nozzle with several motion stages enabling the reactor to create three-dimensional objects by traversing and rotating the substrate.3,4 The control of the plane geometry of the structures was accomplished by programming the substrate stages to move in the desired patterns.

The most typical deposits that have been made in this manner at Georgia Tech have been of pyrolytic carbon; using the reagent gases methane and hydrogen at or below atmospheric pressure. All experiments described here used 75% CH4, 25% H2. The pressure in the reaction zone was held at 1 atm and the reagent nozzle flow was set to 300 sccm CH4 and 100 sccm H2.Layered line growth using the LCVD process is dependant on several environmental variables that can greatly affect the rate and type of deposit that is produced. The most important of these is temperature. Some of the experiments presented below utilized a precision temperature control system. The temperature control was accomplished through use of a thermal imaging camera capable of resolving the temperature in the region of the laser spot to 3.5 µm per pixel. The camera’s software allows for the creation of one or more ‘Regions Of Interest’ (ROIs) within which the average temperature can be calculated.

The camera’s software allows for the creation of one or more ‘Regions Of Interest’ (ROIs) within which the average temperature can be calculated. The average ROI temperature is fed back to a PID controller for the laser power so as to maintain a target temperature as set by the operator. ROIs for these experiments were typically areas of about 0.246 mm2 and temperature set points for the regions were varied between 1600 and 1700°C. Note that the ROI temperature does not necessarily correspond to the substrate temperature, but is a useful means for controlling the LCVD reaction.Another factor in building layered structures is the speed at which the substrate is traversed beneath the laser spot, i.e. the scan speed. Many different speeds have been tested, and the tendency for heat to conduct away through the substrate during the initial stages of a wall enforces an upper limit on the initial scan speed.

A lower limit is also imposed on the scan speed below which deposits may begin growing away from the substrate into fibers.2,5 The fastest speed that has been used thus far with successful initiation under the given conditions has been about 42.3 µm/s and thus was used for these experiments.A final factor in making deposits is the substrate material.6 The Georgia Tech system uses a sacrificial substrate system, and these particular experiments used two graphite disk substrates with different levels of porosity.ResultsExperimentation with deposition of multi-layered carbon structures without a means of automatically controlling temperature and laser power yielded some interesting results that are consistent with the findings of others.2 The procedure was typically to bring the reactor to running conditions and simultaneously begin translating the substrate and initialize the laser to a specific power level.

Sometimes, the laser power was not altered throughout the course of the experiment while other experiments involved some systematic manual reduction in laser power as the wall height and number of layers increased.62The deposits exhibit non-uniform morphology and size. In particular, the laminated walls tend to develop large obtrusions on their ends while the center regions maintain a more uniform height as is shown in Figure 1.600 µm 600 µm600 µm 600 µmFigure 1. Carbon walls made without process control.7The end obtrusions are likely related to several factors. One possible factor is the tendency for the linear translating stage to decelerate upon completion of a pass causing the end of the line to be exposed to higher temperatures.7 A related factor involves the possibility of the stage pausing at the end of a pass causing the same phenomenon.

In order to study the stage deceleration and pause factors, a control program was implemented to turn off the laser prior to the end of linear movement and restart the laser at the same position on the next pass. Figure 2 illustrates the concept of this control program. The implementation of the stage translation program had no significant impact on the end irregularity.Laser OnStage MovementFigure 2. Diagram of stage control program used to study relationship between stage movement and line end irregularities.7Another possible source of the obtrusions is related to heat transfer within the deposit and substrate. Initially, residual heat from the preceding laser pass leads to higher local substrate temperatures and an increased deposition rate.

Initially, residual heat from the preceding laser pass leads to higher local substrate temperatures and an increased deposition rate. As the wall height increases, the nature of the heat conduction path changes such that the heat at the ends of the wall can conduct easily along the wall in only one direction, whereas the heat in the center region of the wall can conduct in two directions along the wall. During the deposition of such laminated structures as discussed above, observations were made through the thermal imaging camera as to the behavior of the temperature within the laser spot. It was observed that the ends of the lines exhibited higher temperatures earlier in the deposition process as is consistent with the shape of the resulting deposits. Active temperature control can eliminate these undesirable end effects.63Also noteworthy in Figure 1 is the appearance of a crevice in the center of the deposit.

Active temperature control can eliminate these undesirable end effects.63Also noteworthy in Figure 1 is the appearance of a crevice in the center of the deposit. This ‘volcano effect’ could be detrimental to the use of LCVD in producing precise geometric objects, but has been shown to be a temperature or power density related phenomena.8 According to Jean, the volcano effect for carbon using a methane precursor appears when maximum deposit zone temperatures exceed 2500°C.7 This effect reiterates the need for a suitable temperature control system in the LCVD process.Two grades of graphite substrates were tested for carbon deposition and the best results have been produced using the higher porosity graphite. It is theorized that low porosity graphite is more conductive and allows the heat generated by the laser spot to conduct away faster leading to longer deposit initiation times.

It is theorized that low porosity graphite is more conductive and allows the heat generated by the laser spot to conduct away faster leading to longer deposit initiation times. Use of porous graphite would seem to encourage the creation of irregularities due to surface roughness; however, there were no apparent morphological differences between the deposits produced on the different substrates.The implementation of automatic temperature control has allowed for significant advances in the ability to deposit uniform multi-layer carbon structures. The setup of the temperature control system used in these experiments was presented above. Figure 3 presents an image from the thermal imaging and control system during wall deposition.Scan Direction560 µmFigure 3. Thermal image captured during deposition with temperature control.7Laminated deposits created using the temperature control system exhibited reduced occurrences of end obtrusions and the volcano effect.

The elimination of the end effects with temperature control is due to the system’s ability to automatically compensate for the increased local temperatures on the line ends by reducing laser power. Experimentation showed that the volcano effect could be eliminated by maintaining a temperature set point below 1700 °C for the given ROI. Figure 4 illustrates a typical laminated carbon wall built using the stated conditions.64240 µm 240 µmFigure 4. Twenty pass wall deposited using temperature control.7The surface morphology of the walls was examined using a scanning electron microscope (SEM). Figure 5a illustrates a view of the top surface of a wall showing a rippled form. Figure 5b depicts the side of a wall. The side appears homogenous yet slightly stratified in accordance with the individual number of wall layers.16 µm16 µm120 µm 120 µm(a)(b)Figure 5.

Figure 5b depicts the side of a wall. The side appears homogenous yet slightly stratified in accordance with the individual number of wall layers.16 µm16 µm120 µm 120 µm(a)(b)Figure 5. Surface appearance of carbon walls.65Examination of MicrostructureInternal analysis of the structures produced in these experiments yielded some interesting findings. Several samples were mounted in epoxy, cut, and polished to allow for examination of the cross section of the walls. Figure 6 is an example of such a cross section.Smaller volcano effectLayer BoundriesMore pronounced volcano effectPossible c rack20 µmFigure 6. Cross-section of multi-layer carbon wall.7One should first note the uniform density of the deposit. Individual layers of the wall can easily be seen in Figure 6. It is believed that some of the less defined black marks in the photo are artifacts of sample preparation such as residue from the polishing operation.

It is believed that some of the less defined black marks in the photo are artifacts of sample preparation such as residue from the polishing operation. The basic form of the deposit has the characteristic volcano shape, which appears to diminish with increasing numbers of layers. The diminishing of the volcano effect with height is evidence that a temperature controlled process can eliminate sizeable interlayer voids. An important feature observed in the cross section are cracks believed to be related to thermal stress between layers of deposit. One proposed method for addressing these cracks involves using a substrate heater to bring the substrate temperature closer to the reaction temperature and reduce the thermal gradients within the deposited structure. A substrate heater has been implemented into the reactor, but experiments to determine its effectiveness have not yet been conducted.Some features of the deposit likely stem from a lack of robustness in the existing LCVD system.

For example, the variation in the horizontal position of the layers as the height of the wall increases could be due to straightness error associated with the linear stage specified to be ±2.5 µm/25 mm. This can be seen in Figure 6, where the layers offset the width of the deposit66by up to 10 µm. The variations could also be caused by lack of rigidity in the mounting design for the stages.Figure 7 is another wall cross section that illustrates the integrity of the multi-layer carbon walls. This sample does not show evidence of the volcano effect. The lower layers of the structure have a more rippled interface than the upper layers which again illustrates the importance of the temperature control process in dampening the effects of local abnormalities. This structure does not have the horizontal oscillations shown in Figure 6.Smooth interfacePolishing artifactsRippled inte rfaceGraphite substrateFigure 7.

This structure does not have the horizontal oscillations shown in Figure 6.Smooth interfacePolishing artifactsRippled inte rfaceGraphite substrateFigure 7. Cross section of wall without any volcano effects.ConclusionsExperiments have shown that a temperature controlled LCVD process is capable of growing multi-layered pyrolytic carbon structures. The propagation of irregularities through the structures is reduced by the control process, resulting in a uniform deposit with a flat top and straight side walls. Viewing of the wall cross sections indicates a generally solid internal structure, free of large voids or defects. The interlayer bonding may be adversely affected by the thermal stresses associated with the laser heating. These thermal stresses could be the source of cracks shown in Figure 6. More testing is required to determine the extent of this apparent cracking and its impact on the functionality of LCVD multi-layered structures.

More testing is required to determine the extent of this apparent cracking and its impact on the functionality of LCVD multi-layered structures. In conclusion, we believe this work to be an important building block for the future creation of more complex 3D layered structures.AcknowledgementsThis work is supported by the National Science Foundation and the Georgia Institute of Technology.67References1. C. Duty, D. Jean, and W. J. Lackey. “Laser Chemical Vapor Deposition: Materials, Modeling, and Process Control,” Accepted for publication in International Materials Reviews, 2001.2. D. Messia, J. Pegna, W. H. Lee, D. DeAngelis. “Layered Micro-Wall Structures from Gas Phase,” Solid Freeform Fabrication Symposium Proceedings, 1997, 231-239.3. D. Jean, C. Duty, B. Fuhrman, W. J. Lackey. “Precision LCVD System Design with Real Time Process Control,” Solid Freeform Fabrication Symposium Proceedings, 59-65, 1999.4. C. Duty, D. Jean, and W. J. Lackey.

A layer of fluoro-silane was then grown on the surface of the mold to promote the release of the mold after imprinting. To test the imprinting capacity of the mold, poly (ethylene glycol) diacrylate was patterned by the mold and the results were analyzed by a scanning electron microscope (SEM).laser was used toinduce TPPI.IntroductionNanoimprinting lithography (NIL) is a low-cost, high-resolution, and parallel nano- fabrication technique that has been intensely studied in recent years. NIL is capable of fabricating nanoscale structures, which, due to diffraction limit, cannot be accomplished the commonly-used by most other parallel photolithography process [1, 2]. Furthermore, because of its parallel fabrication mode, NIL could fabricate samples much faster than could other nanofabrication methods such as Electron Beam Lithography (EBL).

Furthermore, because of its parallel fabrication mode, NIL could fabricate samples much faster than could other nanofabrication methods such as Electron Beam Lithography (EBL). Unfortunately, fabricating nanoimprinting molds usually requires numerous costly and often unavailable cleanroom equipments and time-consuming steps, including EBL and reactive ion etching (RIE) [3].fabrication methodssuch asIn this report, we demonstrate the plausibility of creating a low-cost polymer nanoimprinting mold by using the two-photon polymerization (TPP) technique. TPP fabrication technique has already made sub-100 nm features by breaking the diffraction limit which used to restrain the resolution of photo-polymerization fabrication. However, TPP is an extremely slow serial (point-by-point) fabricating process, which has precluded it from being used in the mass production of nanostructures.

However, TPP is an extremely slow serial (point-by-point) fabricating process, which has precluded it from being used in the mass production of nanostructures. Our project, by using TPP to make reusable nanoimprinting molds, indirectly enables the utilization of TPP in mass production. Moreover, a variety of polymers can be used to make such molds, which ensures that it is relatively easy to obtain the desirable bulk and surface properties for polymer molds suitable for different imprinting projects. Poly (ethylene glycol) diacrylate (PEGDA) was patterned in the study to demonstrate the imprinting capacity of the polymer molds made by TPP.II. Materials and MethodsAcrylic Polymer Preparation for TPP Fabrication. The acrylic polymer we used for the TPP fabrication was dipentaerythritol pentaacrylate (DPPA, Sartomer Inc. Exton, PA), which was used as received.

The acrylic polymer we used for the TPP fabrication was dipentaerythritol pentaacrylate (DPPA, Sartomer Inc. Exton, PA), which was used as received. To increase the TPP rate of DPPA, 1% photoinitiator (Irgacure 819, Ciba Specialty Chemicals, Tarrytown, NY) was thoroughly mixed with DPPA by heating the mixture in 80 °C silicone oil (Fisher Scientific Inc. Pittsburgh, PA).The acrylic polymer should adhere on a glass substrate in order to be used as a mold. To enhance the adhesion with DPPA, the glass substrates were treated for 5 minutes in a 1mM solution of 3-trichlorosilyl propyl methacrylate (TPM, Fluka Chemicals, Milwaukee, WI) in a 4:1 mixture of heptane (Fisher Scientific Inc. Pittsburgh, PA) and carbon tetrachloride (Sigma-Aldrich Inc. Milwaukee, WI). Then the substrates were washed by carbon tetrachloride and dried by nitrogen.

Pittsburgh, PA) and carbon tetrachloride (Sigma-Aldrich Inc. Milwaukee, WI). Then the substrates were washed by carbon tetrachloride and dried by nitrogen. Finally, in order to get rid of the chlorhydric acid generated in the previous step and the residual water on the surface, the substrates were soaked in DI water for 2 minutes and then baked on a hot plate (Super- Nuova, Barnstead Thermolyne Corporation, Dubuque, IA) at 100 °C for 5 minutes.TPP Fabrication Process. A Ti:sapphire femtosecond laser (Vitesse, Coherent Inc. Santa Clara, CA) producing 100-femtosecond duration pulses with a central wavelength of 800 nm at a repetition rate of 80 MHz was used as the excitation source for TPP. To use the full capacity of the large numerical aperture oil-immersion objective lens, the laser beam was expanded by a 2X beam expander to overfill the incidence aperture of the 100X oil-immersion objective lens (N.A. = 1.3, Fluar, Carl Zeiss MicroImaging, Inc.

= 1.3, Fluar, Carl Zeiss MicroImaging, Inc. Thornwood, NY) which focused the laser beam into the volume of DPPA. The laser power we used in the mold fabrication was 5 mW, measured by a power meter (PowerMax 500D, Molectron Detector Inc. Portland, Oregon) before the laser beam entering the objective lens. The mold structures were generated by moving a motorized stage (Newport Inc. Irvine, CA), which held the glass substrate, at a speed of 50 µm / second. The fabrication process was monitored in situ by a charge-coupled device (CCD) camera (CV-S3200, JAI Inc. San Jose, CA). The mold was developed in ethanol (Decon Labs Inc. King of Prussia, PA) for 3 minutes to wash away uncured DPPA after TPP fabrication.Treatment of the Mold after the TPP Fabrication.

San Jose, CA). The mold was developed in ethanol (Decon Labs Inc. King of Prussia, PA) for 3 minutes to wash away uncured DPPA after TPP fabrication.Treatment of the Mold after the TPP Fabrication. To get rid of the unreacted acrylate groups on the mold surface, the mold was soaked in a solution of 1 wt% photoinitiator (Irgacure 2959, Ciba Specialty Chemicals, Tarrytown, NY) in DI water and was irradiated with a UV lamp (OmniCure S2000, Exfo Life Sciences & Industrial Division, Mississauga, Ontario, Canada) for 2 minutes. After that, a fluoro-silane based releasing layer was grown on the mold surface to facilitate detachment of the mold during the imprinting process. Figure 1 shows the schema of the silanization process. First, two glass petri dishes, containing several drops of silicon tetrachloride and the mold, respectively, were placed in a vacuum oven (Isotemp Vacuum Oven Model 280A, Fisher Scientific Inc. Pittsburgh, PA) for 2 hours.

Pittsburgh, PA) for 2 hours. During this process, a layer of silicon tetrachloride (Acros Organics, Morris Plains, NJ) was grown on the mold surface, as Figure 1 (a) illustrates. Then the mold was soaked in DI water for 2 minutes, during which time the chlorineions of silicon tetrachloride were replaced by hydroxyl groups of water molecules, as Figure 1 (b) shows. This replacement greatly increased the number of hydroxyl groups on the mold surface, making it easier for fluoro-silane molecules to bond onto. After that, the mold was baked on a hot plate at 100 °C for 10 minutes to strengthen the newly formed layer of hydroxyl groups by forming -Si-O-Si- bonds between adjacent silicon tetrachloride molecules. Next, a layer of fluoro-silane -- tridecafluoro-1,1,2,2- tetrahydrooctyl-1 trichlorosilane (TFTT, United Chemical Technologies Inc.

Next, a layer of fluoro-silane -- tridecafluoro-1,1,2,2- tetrahydrooctyl-1 trichlorosilane (TFTT, United Chemical Technologies Inc. Bristol, PA) was grown on the mold surface following similar steps as the ones used for growing silicon tetrachloride, with the only difference being that TFTT, instead of silicon tetrachloride, was placed onto one of the two glass petri dishes. Figure 1 (c) illustrates the reactions occurred in this step. Finally, the mold was again soaked in DI water for 2 minutes and backed on a hot plate at 100 °C for 10 minutes. Figure 1 (d) shows the final result. After the fluoro-silane had been successfully grown on the mold, the mold was stored in a vacuum desiccator (Desi-Vac, Fisher Scientific Inc. Pittsburgh, PA).Figure 1.

After the fluoro-silane had been successfully grown on the mold, the mold was stored in a vacuum desiccator (Desi-Vac, Fisher Scientific Inc. Pittsburgh, PA).Figure 1. Schematic diagram of the treatment of the mold surface: silicon (a) tetracholoride reacted with the hydroxyl groups on the mold surface; (b) the mold was soaked in DI water to form more hydroxyl groups on its tridecafluoro- surface; 1,1,2,2-tetrahydrooctyl-1 trichlorosilane reacted with the hydroxyl groups on the mold surface, RF = (CH2)2(CF2)5CF3; (d) a fluoro-silane layer was formed on the mold surface.(c)Imprinting in PEGDA. The imprinting work was conducted in a mask aligner (SUSS MA-6, SuSS Microtech, Munich, Germany), following a procedure outlined below. A more detailed discussion regarding the procedure was described in our previous study [4]. Firstly, PEGDA (MW 258, Sigma-Aldrich Inc.

A more detailed discussion regarding the procedure was described in our previous study [4]. Firstly, PEGDA (MW 258, Sigma-Aldrich Inc. Milwaukee, WI) was thoroughly mixed with 1% Irgacure 2959 through 5-minute sonication (Branson Ultrasonics Corporation, Danbury, CT) to increase its cross-linking rate. Next, the mold was fixed on the sample stage of the mask aligner by two strips of 100 µm-thick sticky tapes. The sample substrate that served as a transparent dummy mask was similarly affixed onto a glass dummy plate by two sticky tapes. After that, the glass dummy plate with the samplesubstrate affixed on it was mounted to the mask holder of the mask aligner by vacuum suction so that the glass substrate was directly above and facing the mold. Lastly, a drop of PEGDA was deposited on the mold surface by a glass pipette. The sample stage was then lifted upwards until the tapes on the mold and on the sample substrate contacted with each other.

The sample stage was then lifted upwards until the tapes on the mold and on the sample substrate contacted with each other. The PEGDA was allowed to spread for 3 minutes and was then irradiated by 365 nm UV light for 30 seconds. After the exposure, the mold was detached from the sample substrate by the automatic descent of the sample stage. After the imprinting process, the mold was taken from the sample stage and could be used in subsequent imprinting processes without any further surface treatment.III. Results and DiscussionThe mold and the imprinted PEGDA structures were examined by a scanning electron microscope (SEM, Supra 40VP, Zeiss, Carl Zeiss SMT Inc., Peabody, MA). As figure 2 (a) shows. The width and height of each line on the mold is 400 nm and 1.2 µm, respectively, and the distance between adjacent lines is 5 µm. Figure 2 (b) presents a side view of the parallel lines and provides a direct view of the height of the lines.

Figure 2 (b) presents a side view of the parallel lines and provides a direct view of the height of the lines. Figure 2 (c) and (d) show that the imprinted PEGDA structures consist of parallel trenches of the same dimensions and spacing as the parallel lines on the mold, thus demonstrating that our imprinting process accurately transferred the patterns on the DPPA mold to PEGDA over a large area. While the widths and the spacing of the trenches can be easily measured from the top-view SEM image on figure 2 (c), the height of the trench cannot be obtained that way. It cannot be obtained by an atomic force microscope (AFM) scanning either because the trench is too narrow for the AFM tip to approach. Therefore, to measure the height of the trenches, the cross-sectional view of the trenches, figure 2(d), was needed.Figure 2.

Therefore, to measure the height of the trenches, the cross-sectional view of the trenches, figure 2(d), was needed.Figure 2. SEM images of the imprinted PEGDA structures and the DPPA mold after imprinting: (a) DPPA mold of 400 nm- wide lines with a period of 5 µm; (b) side view of (c) the DPPA mold; imprinted PEGDA structures of 400nm wide trenches with a 5µm period; (d) cross- sectional view of the PEGDA trenches, which perpendicularly were cut to expose the cross- sections.Our work demonstrates that imprinting molds made by TPP could successfully transfer nanoscale patterns to PEGDA. To successfully transfer the patterns, the mold must be easily detached from the imprinted materials after imprinting. Thus it was crucial to grow a layer of silicon tetrachloride on the mold surface before growing the releasing layer of fluoro-silane.

Thus it was crucial to grow a layer of silicon tetrachloride on the mold surface before growing the releasing layer of fluoro-silane. This is because silicon tetrachloride tripled the number of hydroxyl groups on the mold surface so that a much denser layer of fluoro-silane was able to bond onto it. Another method to increase the number of hydroxyl groups on the polymer surface is by transforming some methyl groups of the polymer to hydroxyl groups via RIE [5]. This technique is a promising alternative treatment for the mold surface since it is simpler than the silicon tetrachloride treatment used in our experiment. is also noteworthy that some fluoro-based polymers such as perfluoropolyether (PFPE) have intrinsic ultralow surface energy, as a result of which molds made of such materials have inherent anti-attaching ability so they can imprint without any surface treatment.

Previous studies have demonstrated the superior imprinting capability of PFPE nanoimprinting molds made from a master silicon mold by UV polymerization [6]. We think PFPE can also be patterned by TPP to make nanoimprinting molds, which will further simplify our fabrication processes.ItIV. ConclusionsAn acrylic polymer nanoimprinting mold was generated by TPP fabrication using a femtosecond laser. The mold was treated by silanization steps to promote detachment of the mold from the imprinted materials. The SEM images of the mold and the imprinted PEGDA patterns show that a group of 400 nm-wide parallel lines were successfully transferred to PEGDA, which demonstrates the imprinting capacity of the mold. The fabrication process of the acrylic polymer mold was simpler than that of the traditional silicon/quartz mold. We boldly expect that the TPP technique will be widely used in making nanoimprinting molds in future.V.

brication of Parts Containing Small Features using StereolithographyJouni P. Partanen3D Systems, 26081 Avenue Hall, Valencia, California 91355ABSTRACTThe StereoLithography (SL) process has benefited from many advances in the last four to five years. These include new resins with reduced shrinkage and curl distortion, enhanced software, and improved scanning techniques. One can produce highly accurate parts for most dimensions within a few mils of the design value as shown in numerous accuracy and benchmarking studies. SLA systems use a laser beam focused to a spot size of 200 -250 ~m. This limits the range of applications where SL can provide accurate models to parts which do not contain very small features i. e. wall thickness values less than about 300 ~m. Industries that manufacture products involving components with small features include electronics and medical. In this presentation we describe an extension of the SL technology to applications involving small features.

In this presentation we describe an extension of the SL technology to applications involving small features. This capability is achieved by reducing the laser focal spot size in an SLA-250 to 75 ~m. The technological principle behind the spot size reduction is described in the presentation, together with process issues and applications of the technology.1. INTRODUCTIONCommon applications of Rapid Prototyping and Manufacturing (RP&M) are in the automotive and consumer electronics industries. Typical models build by RP&M techniques are from a few centimeters to half a meter in all three dimensions[1,2]. Even if these applications often require parts to be accurate within 0.1 mm or better of its design dimensions, they do not usually have features with dimensions in the tenth of a millimeter scale. Thus most of the RP&M techniques have not been developed to be capable of 100 micron or better resolution.

Thus most of the RP&M techniques have not been developed to be capable of 100 micron or better resolution. However, there are many applications of RP&M in the electronics and medical industries which would require less than two hundred micron resolution. One example is electronic circuit board connectors that can have pin densities as high as 2 pins per millimeters.This paper describes an extension of a commercial StereoLithography Apparatus (SLA) to applications requiring about 100 micron resolution. A specialized unit for ultra high resolution (about 10 micron) is described in Ref. [3]. In section 2 we discuss what is required from an RP&M system to achieve a certain resolution. Section 3 describes how the resolution of laser based RP&M is controlled by diffraction. Some properties of 3D Systems Stereolithography Apparatus, SLA-250, are presented in the following section together with a solution for increased resolution.

Some properties of 3D Systems Stereolithography Apparatus, SLA-250, are presented in the following section together with a solution for increased resolution. The performance of the high resolution SLA is characterized63using a generic connector test part as described in Section 5. Some real world applications are shown in the following section.2. RP&M EQUIPMENT RESOLUTIONAll the various RP&M systems build 3 dimensional parts in layers. The resolution in the direction normal to the layers (defined here as the z direction) is limited by the layer thickness. The resolution in the plane of the layers (x-y plane) depends on the method used to define the borders of each layer. In RP&M systems which use lasers to shape individual layers, the laser spot size at the working medium is one of the main factors defining resolution. Figure 1 shows schematically the effect of the spot size on the smallest feature that can be built by two different approaches.

Figure 1 shows schematically the effect of the spot size on the smallest feature that can be built by two different approaches. In the case where a laser is used to solidify the either liquid or powder into the desired shape the minimum positive feature is controlled by the laser spot size (Fig. lea)). If the laser is used to remove material around the part the spot size controls the minimum negative feature (Fig. 1(b)). Because useful parts contain both positive and negative features we can say that the resolution is essentially limited by the spot size. In those RP&M methods which do not use lasers the circle in Fig. 1 can be thought to represent the droplet size, the width of extruded material, or the width of the cutting tool.(a)(b)Laser SpotSmallest positive featureSmallest negative featureLaser SpotFig. 1.

1. Relation between the smallest features that can be build and the laser spot size for laser based RP&M technologies in the two cases where laser is used either to solidify (a) or remove (b) material.64Even if the spot size in Fig. 1 is the limiting factor for system resolution it is not the only factor affecting the x-y resolution. There are number of other effects which have to be taken into account: 1) The solidification process can extend beyond the photoiniation point. 2) The effect of heating can extend beyond the spot size in both sintering and cutting. 3) The particle size and the size of nonuniformities in the raw material affect the resolution. 4) The small features have to be sufficiently strong to survive the process as well as their own weight. The aim of this paper is to demonstrate that the laser spot size is the main factor defining the resolution of SL, and features as small as 100 Jim can be built using SL, when the spot size is reduced to 75 Jim.3.

LASER SPOT SIZEThe physical principle which controls the spot size at the working surface of an RP&M It relates the spot radius Wo at the working surface, (measured at lie 2 of the distancef from the working surface to the last the diameter D of the beam at that element, and the laser beam qualitysystem is diffraction. the peak intensity), to the laser wavelength /l" optical element, characterized by the mode purity parameter M 2The relation can be written as:where the numerical value of the coefficient C depends on the exact definition of the quantities given above, as well as the precise mode shape[4,5]. However, for the present purpose it is sufficient to know that C is close to 1.5.The equation (1) clearly demonstrates the benefits of short wavelength lasers for high resolution RP&M technology.

However, for the present purpose it is sufficient to know that C is close to 1.5.The equation (1) clearly demonstrates the benefits of short wavelength lasers for high resolution RP&M technology. lasers used in stereolithography have a wavelength /l, from 0.32 - 0.36 Jim, about 1/30 of the wavelength /l" 10.6 Jlm, of the carbon dioxide laser commonly used in other RP&M techniques.The ultraviolet4. HIGH RESOLUTION SLA-250The spot diameter 2Wo of a commercially available SLA-250 is typically 200 - 250 Jlm. The three parameters in Eq. (2) which we can use to reduce the laser spot diameter in an SLA 250 are: the distance ffrom resin surface to the scanning mirror (last optical element), the mirror diameter D, and the mode purity parameter M 2 • Changing either the size of the scanning mirrors to allow a larger beam diameter D, or the distance f from the scanning mirrors to the resin surface, would involve a significant hardware development program.

The mode purity parameter M 2 of an HeCd laser used in SLA-250 is typically about 5. It can be reduced to about 1.5 (laser beam clean-up) with a rather simple optical arrangement, allowing a spot diameter 2Wo of about 75 Ilm at the resin surface.In addition to the reduction of the spot size, the laser power will also diminish in the laser beam clean-up process. The reduction of laser power, far from being a negative, is actually65(1)useful in order to preserve drawing accuracy with the smaller spot size. This can be seen fromthe expressionb" the laser scanning speed Vswhere Ee and Opare resin parameters (critical exposure and penetration depth, respectively). For constant cure ,depth Cd, the laser scanning speed Vs is proportional to the ratio of the laser power PL and the spot radius Woo Thus, in order to avoid excessively high scanning speeds, the laser In some cases, even further reduction power should be reduced in proportion with the spot size.

- 400 micron feature~_-,O!!::-_--+,j0050100150200250Cure depth (microns)Fig. 4. Experimentally determined line width Lw as a function of the border cure depth Cd inSmall Spot SLA.Small Spot SLA is being used presently by five companies: One of these is applying the technology to the medical field in developing smaller tools for less invasive surgery. Another company is using their Small Spot SLA in the electrical connector industry as described above. The other organizations are applying the high resolution capability in many different fields as a internal job shop for a multi-industry company, and independent service bureaus. There are also companies using other approaches to reduce the spot size in stereolithography to a diameter between 100 ~m and 200 Jlm.7. CONCLUSIONWe demonstrate in this paper that stereolithography can be easily extended to higher resolution by simply reducing the spot size of the laser on the working surface of an SLA system.

CONCLUSIONWe demonstrate in this paper that stereolithography can be easily extended to higher resolution by simply reducing the spot size of the laser on the working surface of an SLA system. We have successfully built connector test parts that have 300 ~m pitch (3.3 pins per mm).69[1][2][3][4][5]REFERENCESP. F. Stereolithography," Society of Manufacturing Engineers, Dearborn, Michigan, 1992.Jacobs, Ed.,"Rapid Prototyping & Manufacturing: Fundamentals ofP. F. Jacobs, Ed., "Stereolithography and other RP&M Techniques," Society of Manufacturing Engineers, Dearborn, Michigan, 1996.K. Ikuta, K. Hirowatari, and T. Ogata, "Ultra High Resolution Stereolithography for Three Dimensional Microfabrication," Proceedings of the Fifth International Conference on Rapid Prototyping - 1994, University of Dayton, Dayton, Ohio, 1994.'A. E. Siegman, Lasers, University Science Books, Mill Valley, CA, 1986.W. F.

brication of PEM Fuel Cell Bipolar Plates by Indirect SLS Ssuwei Chen, David L. Bourell, Kristin L. Wood Department of Mechanical Engineering, Laboratory for Freeform Fabrication The University of Texas at Austin, Austin, Texas, 78712Reviewed, accepted August 23, 2004AbstractThe paper presents a new manufacturing technique involving Selective LaserSintering (SLS) for proton exchange membrane fuel cell (PEMFC) bipolar platefabrication. A material system for bipolar plate fabrication was identified to satisfy boththe cell performance requirement and SLS operation restriction. Carbonization and liquidepoxy infiltration are subsequently performed following the completion of SLS greenbipolar plate. The finished SLS bipolar plate showed impressive surface finish andmechanical strength, and a single fuel cell was assembled with two SLS end plates andmembrane electrode assembly (MEA) in between. Various physical property tests wereperformed with positive results.

Various physical property tests wereperformed with positive results. Fuel cell performance (voltage vs. current density,voltage vs. time, etc.) will be assessed in the near future.1. IntroductionBipolar plates, most fabricated by machining of pure graphite, are one of the mostexpensive components in proton exchange membrane fuel cells (PEMFC). Extreme caremust be taken during the costly and time-consuming machining process, mainly becauseof the brittleness of graphite. The layer-based nature of selective laser sintering (SLS)offers several advantages for bipolar plate development and manufacturing. This additiveprocess possesses the ability to manufacture complex geometries that are otherwisedifficult to obtain, among which are the grooved flow fields on both sides of bipolarplates.

This additiveprocess possesses the ability to manufacture complex geometries that are otherwisedifficult to obtain, among which are the grooved flow fields on both sides of bipolarplates. Furthermore, integrated cooling channels reserved for improving thermalmanagement, can be simultaneously formed on the central surfaces of the plate ratherthan by gluing two half plates back-to-back as is used in many other techniques. Eventhough application of SLS for bipolar plate fabrication may not be suitable for massvolume production at present, it will significantly benefit research in fuel cell flow fielddesign since prototypes can be done in a much shorter period of time compared to244traditional machining.2. Requirements for Bipolar PlateAn optimized bipolar plate has to fulfill several requirements to maximize cellperformance.

Requirements for Bipolar PlateAn optimized bipolar plate has to fulfill several requirements to maximize cellperformance. From the material point of view, high electric conductivity could minimizethe ohmic losses, meaning that voltage drop across the bipolar plate can be reduced. Eventhough it is always desirable to obtain plate materials with low bulk resistivity so as toreduce the voltage loss, it has been shown that the potential drop across theplate/electrode interface is many orders of magnitude more than that across the bipolarplate itself [1]. Therefore, finding a material more electrically compatible with theelectrode material plays a more important role from the viewpoint of minimizing overallpotential losses.Bipolar plates also require very low gas permeability with respect to reactantgases.

Since hydrogen and air/oxygen are flowing in the flow fields on opposite sides ofthe bipolar plate, it is important to make sure that both gases are not mixed by permeationthrough the plate. Gas tightness of the bipolar plate not only prevents the cross leaks ofreactant gases, but also removes the possibility of dangerous gas leakage into the exteriorfuel cell environment. As far as safety and fuel saving are concerned, bipolar plates witha high gas impermeability are a top priority.Both heat and water/vapor are byproducts of the PEM fuel cell net reaction.Bipolar plate materials having a high thermal conductivity are desired to make best use ofthe produced heat. The overall fuel cell efficiency could be boosted up to 70% [2] if mostof the heat can be cycled. With high thermal conductivity, the bipolar plate is able totransfer heat efficiently to the designed heat exchange unit and maintain the systemtemperature in the operating range.

With high thermal conductivity, the bipolar plate is able totransfer heat efficiently to the designed heat exchange unit and maintain the systemtemperature in the operating range. This is especially important in automotiveapplications since heat produced can be treated as the source of energy for the heatingunits. Additionally, high heat conductivity also ensures low temperature gradients inbipolar plate and eliminates the possible thermal strains and stresses.Corrosion resistance and chemical stability are the two other properties importantfor bipolar plates. PEM fuel cell bipolar plates are generally in contact with an acidicMEA and operate in a hot and humid environment [3]. Without a certain level ofresistance to corrosion, a bipolar plate would oxidize and corrode easily during its long-term lifetime.

Without a certain level ofresistance to corrosion, a bipolar plate would oxidize and corrode easily during its long-term lifetime. Corroded bipolar plates can no longer provide the standard mechanical245strength and may form contamination in the MEA membrane, resulting in the degradationof fuel cell performance [4].In addition, bipolar plates need to be slim and light to keep the stack volume andweightlow. Low density plate materials satisfytheserequirements. Manyelectrochemical parameters, such as volumetric power, volumetric energy andgravimetric capacity, etc., can be increased dramatically by using low specific weightmaterials. Finally, the cost of both bipolar plate material and manufacturing techniqueneeds to be lowered to promote commercialization of PEM fuel cells. PEM fuel cellscurrently are prohibitively expensive, due significantly to bipolar plate cost [5].

PEM fuel cellscurrently are prohibitively expensive, due significantly to bipolar plate cost [5]. Thus,searching for an inexpensive bipolar plate material with the capability for easy fabricationhas become the major task for industries that are trying desperately to make commercialPEM fuel cells.3. Material Selection and Manufacturing Procedures3.1 Material Selection and Screening TestA material system was chosen consisting of graphite as the matrix material andphenolic as the binder. Phenolic is well known for giving excellent bonding strength uponcuring and more importantly, a higher carbon yield can be obtained when completecarbonization is performed [6].

Phenolic is well known for giving excellent bonding strength uponcuring and more importantly, a higher carbon yield can be obtained when completecarbonization is performed [6]. Table 1 lists a comparison of carbon yield from someselected hydrocarbon polymers.MaterialTheoretical carbon Content (%) <80 Phenolic resin <75 Furan resin Epoxy resin <75 Unsaturated polyester resin <60Carbon yield (%) after Pyrolysis 55-70 50-60 20-30 15-25Table 1 Carbon yield for selected materialsThe nature of graphite makes it an excellent choice for bipolar plate fabrication asfar as material requirements are concerned.

However, extremely poor wettability andsurface roughness render it relatively unfit for processing using powder metallurgy.Wettability of graphite powder plays a very important role in this case since phenolicresin, which serves as the binder, is expected to be fused into a preform in the liquidphase to provide the initial binding strength between graphite particles during sintering.246The strategy to overcome the low wettability barrier is to mix finer particle size phenolicresins with relatively coarser pure graphite particles. It is expected that a large amount ofphenolic resin powder will become uniformly dispersed over the larger graphite particlesurfaces, which will facilitate better binder flow during sintering and promote the bindingstrength. Increased binder level is advantageous to increased green strength; however,more geometric shrinkage occurs accompanying this. There should be a compromise asthese two issues are considered.

Increased binder level is advantageous to increased green strength; however,more geometric shrinkage occurs accompanying this. There should be a compromise asthese two issues are considered. Figure 1 shows a mixture of phenolic and graphitepowders in which finer phenolic resins are scattered over the coarser graphite powders.Packing of powderis closelyrelated to the green part density andhence, green strength. Particle size andparticle shape are the two dominantfactorsinfluencingthe efficiency ofpowder packing [7]. The amount offriction of a powder increases as itssurface areaincreases. More frictionFig.1 Roller mixed mixture of graphitebetween particles results in less efficientand phenolic powdersflow and packing [8]. As shown in Figure2, graphite particles are quite irregularand flaky in shape and rough on the surface, which make it hard to be compactedacceptably.

As shown in Figure2, graphite particles are quite irregularand flaky in shape and rough on the surface, which make it hard to be compactedacceptably. One way to reduce the friction between particles is to use graphite powderscoarser in size since most fine powders will not flow because of their high inter-particlefriction. Adjustment of particle size distribution may be beneficial for more compactpacking of powders as well. Packing efficiency can be improved by mixing selectivesizes so as to fill interstices between particles.247However,too wideadistributionofparticlesizeincreases the net surface area sothat poor packing results. Theoptimized size distribution wasinitially screened for better powderflow and surface finishing of theparts. Three groups of graphitepowderswereobtainedFig. 2 SEM of pure graphite particlescommercially and some of theirphysical properties are summarizedin Table 2.

2 SEM of pure graphite particlescommercially and some of theirphysical properties are summarizedin Table 2. The BET surface area was measured by a Micromeritics ASAP 2010 SurfaceArea Analyzer.Source of graphite powder GrafTech (GP44R) 2.25Theoretical Density ( g/cm3)GrafTech (GS150E) Alfa Aesar (Stock#43816)2.242.25Apparent Density ( g/cm3) 0.2540.35---Size Distribution90% < 41 (cid:181)m 50% < 18 (cid:181)m 90% < 100 (cid:181)m 50% < 55 (cid:181)m50~140 meshBET Surface Area ( m2/g)11.354.031.25( 105~297 (cid:181)m)Table 2 Properties of graphite powder examinedA screening test was executed with a high temperature furnace to make bar-shapesamples and their mechanical strength and surface smoothness were evaluated.Eventually, GrafTech GS150E graphite powder was chosen for the material system witha combination of at least 25% to 30% phenolic resin powder obtained from GeorgiaPacific Resins, Inc.

The phenolic powder was 31 (cid:181)m initially but was ground in HJECompany, Inc. to a final size of 11 (cid:181)m.2483.2 SLS of Green Bipolar PlateThe roller mixed graphite and phenolic powder were subsequently loaded into aSinterStation 2000 SLS machine to build green bipolar plates. Curling took place duringgreen part formation. This unwanted effect can be mitigated through bulk powder bedtemperature control. In the direct SLS processes, powder bed temperature is usuallyraised up to a point close to the fusing point of sintering material so as to minimize thethermal gradient [9, 10]. However, in indirect SLS processes involving multiple materialconstituents, this strategy is difficult to fulfill especially for graphite having high meltingpoint and a thermosetting material, phenolic. Figure 3 shows the DSC curve for phenolicresins.

Figure 3 shows the DSC curve for phenolicresins. A powder bed temperature of 55”C was chosen since almost no criticalthermodynamic events took place until around this point during the irreversible curingprocess. The first endothermic peak corresponds to melting of novolac resin while thesecond broad peak corresponds to heat of solution of hexamine in the molten resin [11].The only exothermic peak in the graph represents the cross-linking mechanism ofnovolac resin and hexamine, which occurs at around 150”C. In addition, a laser power of14 watts, a laser scan spacing of 0.003 inch and a layer thickness of 0.004 inch were usedto produce optimum plate quality.25)W W m m ( ( w w o o F l t F a e H t a e H)l201510Novolac Resin MeltHexamine Resin MeltCrosslinking Mechanism5SLS Powder Bed Temperature005055100150200Temperature (C)Fig.

3 DSC curve for phenolic resin2492503.3 Carbonization and Post ProcessingParts produced from SLS are expected to be more porous, weaker and lower inconductivity compared to machined bipolar plates from bulk graphite. Post processingthen would be required to densify the part and attain gas impermeability. The finishedgreen bipolar plate was removed from the powder cake, and loose powders were brushedoff before loading it into a high temperature, vacuum furnace and converting the phenolicresin into carbon residue [12]. The carbonized phenolic resin exhibits an amorphous,glass-like structure as can be verified by XRD pattern shown in Figure 4. The curingprocess was done in a reducing environment by filling the vacuum chamber with eitherargon or nitrogen forming gas to ensure that a maximum amount of carbon residue wasleft during the carbonization.

To prevent curling of brown parts during carbonization, thebipolar plates were all sandwiched by two pieces of graphite slabs during curing.However, some dimensional reduction became more visible in the vertical z-directionsince the weight of graphite slabs made the parts even more compact.9000080000) s t n u o C( y t i s n e t n7000060000500004000030000I200001000000204060Theta (Degree)80100Fig. 4(a) Xray diffraction pattern of GS 150E graphite powder250700600) s t n u o C500400(y t i s n e t n300200I10000204060Theta (Degree)80100Fig. 4(b) Xray diffraction pattern of carbonized phenolic resin, revealing the amorphous crystal structureUpon completion of the curing cycle, the cured brown bipolar plates were cooledand immersed in liquid epoxy resin for a period of a couple of minutes, depending on theviscosity of the resin used [13], to make the porous structure gas tight.

The liquid epoxyresin intruded into open pores of the brown part without producing noticeabledimensional change to the part. Solvents, such as toluene or xylene, were required todilute the as-received liquid epoxy resin since most are too viscous to penetrate into thegraphite pore structure. During infiltration, gas bubbles kept coming out of the partsurfaces, indicating the porous structures were being filled by epoxy resin due to thecapillary forces. Infiltration of at least two times was needed to create a completely gastight plate surface. Infiltrated bipolar plates were subsequently oven dried at 60” C tospeed up the curing process. Figure 5 shows SEM micrographs of the plate surface takenfrom different stages during fabrication procedures.251Fig. 5(a) Green part micrograph after SLSFig. 5(b) Brown part micrograph after phenolic burnoutFigure 5(a) presents the microstructure ofgreen part.

5(a) Green part micrograph after SLSFig. 5(b) Brown part micrograph after phenolic burnoutFigure 5(a) presents the microstructure ofgreen part. As can be seen, the phenolicresin provides bonding between graphitepowders with significant porosity. Thisbonding could be adjusted through binderlevel optimization to give the adequategreen strength for basic plate handlingFig.5(c) Infiltrated part micrographduring post processing. Figure 5(b)demonstratesthe cured brown partstructure details. More open and interconnected pores can be observed due to theburning-off of the phenolic resin, and most of the resin residue was converted intocarbon. The curing process at this stage boosted the electrical conductivity dramaticallyas well as mechanical strength compared to the green part, which may partially beexplained by the reduction of porosity in the part [14, 15].

In Figure 5(c), after epoxyinfiltration, most of the pores were filled up, giving rise to a much more hermetic surfacefinish. Moreover, the infiltrated parts were not inferior to brown parts in electricalconductivity based on two point probe ohmic testing, indicating that the overall carbonphases, inclusive of graphite and converted phenolic resin, are interconnected to eachother without interruption by the insulating epoxy resin.2524. Electrical and Physical Property TestingBulk electric conductivity was measured by four point probe technique, which iscommonly used because of its ability to eliminate the contact resistance between theprobe and test sample. Flexural strength was obtained by three point bend testing (ASTMD790) on a rectangular bar having exactly the same composition as the finished bipolarplate. A Quantachrome ultrapycnometer was used to measure the density of the plate andcorrosion rate was measured electrochemically by a potentiostat.

A Quantachrome ultrapycnometer was used to measure the density of the plate andcorrosion rate was measured electrochemically by a potentiostat. A 0.01 M HCl + 0.01 MNa2SO4 solution was used in corrosion test to mimic the real fuel cell operatingenvironment [3]. Lastly, a Varian Portatest helium gas leak detector was used to evaluatethe gas impermeability. These properties are summarized and compared with someselected novel bipolar plates in Table 3.

Lastly, a Varian Portatest helium gas leak detector was used to evaluatethe gas impermeability. These properties are summarized and compared with someselected novel bipolar plates in Table 3. A single PEM fuel cell assembled by two SLSend plates along with one end plate are shown in Figure 6(a) and two 3 mm thick SLSbipolar plates are shown in Figure 6(b).TypePropertyDepartment of Energy (DOE)Target ValueIGT (molded graphite composite bipolar plate) [16]Oak Ridge National Lab/ Porvair Fuel Cell Technology (slurry molding of carbon/carbon composite bipolar plate) [17, 18, 19] 4200SLS Bipolar PlateFlexural Strength(psi) Electric Conductivity(S/cm) Specific Weight (g/cm3) Corrosion Rate ((cid:181)A/cm2) Gas Permeability (cm3/cm2-s)> 600>100---< 16<16x10-6>3000250~350---< 5< 2x10-6200~3001~1.36< 2x10-6>1730801.27< 95x10-6Table 3 Property of SLS and selected molded bipolar plates253Fig.

6(a) One end plate on the left and a stack assembly on the rightFig. 6(b) Two 97x97x3 mm3 bipolar plates with 8.5 cm2 active area each5. Discussion and ConclusionThe work done so far shows the feasibility of application of SLS to fuel cell bipolar platefabrication. The properties of SLS graphite composite plates are promising and both thematerial and process costs are reduced compared to traditional machining of puregraphite. However, the dimensional shrinkage during brown part formation posed achallenge as far as geometric accuracy is concerned. Fortunately, the dimensionalreduction can be compensated by scaling up the initial plate design since the shrinkageseems to be linear if the powders are well mixed and no density gradient exists in thebatch. Finally, single fuel cell polarization curve and contact resistance between bipolarplate and MEA will be measured to justify the cell performance.

Available athttp://www.eere.energy.gov/hydrogenandfuelcells/pdfs/ive7_haack.pdf255[19] Porvair Fuel Cell Technology. Available at http://www.eere.energy.gov/hydrogenandfuelcells/pdfs/review04/fc_19_haack_04.pdfAcknowledgmentThis research was funded by State of Texas Technology Development and TransferGrant Number 003658 (2001) and by the Office of Naval Research Grant NumberN00014-00-1-0334.256.

brication of Prosthetic Sockets by Selective Laser SinteringW.E. Rogers*, R.H. Crawford**,J.J. Beaman** , and N.E. Walsh* *The University of Texas Health Science Center at San Antonio **The University of Texas at AustinAbstractSolid freeform fabrication technologies offer exciting possibilities for improving product quality by direct manufacture of products. .One example of such product improvement is the fabrication of artificial limb sockets by selective laser sintering (SLS). Currently these sockets are produced at the University of Texas Health Science Center at San Antonio by digitizing the residual limb with a 3D laser scanner, modifying this geometry appropriately using a proprietary CAD system, producing a mold with a computer-controlled milling machine, and vacuum forming the final product.

This paper describes a new manufacturing technique whereby the digital socket data from the CAD system provide input to a SLS workstation to produce the final socket directly, without the intermediate step of fabricating a mold pattern. The advantages of this process include integration of the prosthesis attachment fitting and socket as one component and greater control of local socket geometry for superior stress relief characteristics.Keywordsamputees, computer assisted design, computer assisted manufacture, prosthetics, solidfreeform fabricationIntroductionIn the United States alone, there are more than 400,000 living limb amputees with approximately 60,000 new ones joining the ranks every year. (1) Ninety-seven percent of all these people are candidates for ambulation and could be fitted with prostheses (artificial limbs) if funding were available. The conventional procedure for fabricating and fitting the prostheses is more of an art than a science.

The conventional procedure for fabricating and fitting the prostheses is more of an art than a science. It is non-repeatable and is extremely labor intensive. As pointed out by Klasson (2), there are so many amputees in the world today and many new ones joining with all the regional wars that are being fought that it would take approximately 70,000 new well trained prosthetist to take care of the present needs. It has been suggested by Foort (3)(4) that the only possible way of providing these hundreds of thousands of amputees with appropriate prostheses is to teach machines how to manufacture artificial limbs.The most important aspect of a lower extremity prosthesis is socket design. The socket is the interface between the human and the mechanical support system. Ultimately, the design and fit of the socket is what determines patient acceptance, comfort, suspension and energy expenditure (5).

Ultimately, the design and fit of the socket is what determines patient acceptance, comfort, suspension and energy expenditure (5). All of these factors in unison determine the real utility of the final product.The most common method for designing sockets requires a skilled prosthetist and time consuming manual steps (6)(7). First, the patient's residual limb is wrapped with plaster to acquire the shape. The plaster wrap is used to create a plaster positive. The prosthetist modifies the plaster positive of the residual limb to make a biomechanically correct pattern for a socket. The socket is molded over the pattern. It is necessary to destroy the pattern in order to remove it from the socket. Making a new socket involves repeating all the steps since the pattern is destroyed.158Computer Assisted Design (CAD) and Computer Assisted Manufacture (CAM) techniques are beginning to be used to design and manufacture sockets.

Either a mechanical digitizer or a non-contact laser scanner (8) inputs the residual limb shape into the computer. The prosthetist then uses specialized software to produce a biomechanically correct socket from the limb shape. A numerically controlled milling machine carves the pattern for the socket from a plaster blank. The socket is then made using conventional molding techniques.In all types of socketst the remainder of the prosthetic limb must be attached to the socket. A fitting is bolted through the distal end of the socket or is attached with adhesive to the outside. (Fig 1.) This is a point of frequent structural failure. The other components of the prosthesis are attached to this fitting. The attachment point is somewhat arbitrary and cannot be exactly reproducedt even with current prosthetic CAD software.

The other components of the prosthesis are attached to this fitting. The attachment point is somewhat arbitrary and cannot be exactly reproducedt even with current prosthetic CAD software. The socket design is not integrated with the design of the complete,..-..:-ro;;.:s::;th;;.;e:;,;tl:;:·c...:li=·m::b::.:.:.....--, .All prosthetic CAD/CAM systems use a milling machine to produce a positive mold for a socket. The actual socket is then fonned over the positive. Additional modifications need to be made to the positive mold to allow for attaching a fitting to the bottom of the socket. The fitting is used to connect the socket to the rest of the prosthetic limb. If no modifications are made for a bolt on fitting t then a fitting is attached with adhesive to the bottom of the socket..Below the Knee Prosthettc LimbIt would be ideal to produce the socket directly and avoid producing a plaster positive. Attachment points could be molded as part of the socket.

Attachment points could be molded as part of the socket. Areas of the socket could be selectively strengthened to provide additional support. The research reported here is an investigation into the use of solid freeform fabrication technologies to bypass the molding process by directly producing sockets with integrated attachments.Pylon __.....Prosthetic --a.." FootL..-F..;;ig;..._1......Introduction to SLSSolid Freeform Fabrication (SFF) technologies address the problem of creating three dimensional objects directly from a geometric database without specific tooling or human intervention. The research in SFF at UT Austin has concentrated on a particular process called Selective Laser Sintering (SLS). (9)(10) In SLSt components are built by material addition rather than by material removal by using a directed energy beam which causes individual particles to adhere in selected regions of space. The process begins by first depositing a thin layer of powder into a container.

The process begins by first depositing a thin layer of powder into a container. The powder surface is raster-scanned with a high power energy beamt such as a laser or electron beam. The intensity of the beam is modulated to sinter the powder in areas to be occupied by the part at that particular cross-section. In areas not sinteredt the powder remains loose and may be removed once the part is completed. Successive layers of powder are then deposited and sintered until the entire part is complete.Selective Laser Sintering has the potential to produce accuratet structurally sound three dimensional solid versions of objects designed with a computer and to make such objects available in minutes or hours to the designer. The technology has been licensed commercially to DTM Corporation of AustintTXt to produce parts composed of polymer159powders.

The technology has been licensed commercially to DTM Corporation of AustintTXt to produce parts composed of polymer159powders. Ongoing research at UT Austin is concerned with investigation of SLS of ceramic and metallic powders as well.There are essentially no limitations on the shape of parts that can be fabricated by SLS. Geometric processing consists of two basic steps: slicing and rasterizing. The slicing operation is performed by intersecting the geometric description of the part with planes which represent each of the powder layers. The result is a series of contours which describe the part boundaries in each layer. Rasterization is the process of converting these layer contours into a series of laser toggle points. For this operation, each layer is divided into a number of scan lines. Each scan line is intersected with the contour boundary to produce the toggle points in a manner analogous to scan conversion algorithms used in computer graphics.

Each scan line is intersected with the contour boundary to produce the toggle points in a manner analogous to scan conversion algorithms used in computer graphics. (11) Where the scan line intersects the geometric description of the part the laser is alternately toggled on and off. The toggle points are then passed to the SLS process controller, along with other process parameters, and the part is produced.Part accuracy in the SLS process is affected by the resolution of the geometry processing operations. Currently, parts are made with layer thicknesses of 5 mils. Typical scan line resolution within layers is 20 mils in both directions, although new process software is being developed which will increase the scanning resolution.Prosthetic CAD/CAM at The Health Science Center atThe Departments of Rehabilitation Medicine and Radiology have developed a system for the computer aided design and manufacture of prosthetic sockets for below the knee amputees.

The steps involved are the scanning of the shape of the residual limb of the amputee, the design of the socket from that shape, and the machining of a pattern to mold a socket over. A laser scanner and software to create sockets from limb shapes have been developed at the Health Science Center.In this computerized system, a laser scanner is used to sense the shape of the residual limb. The scanner uses a laser to project a line on the limb and uses two video cameras to determine the location of the projected line in space. The laser and video cameras rotate on a turntable in order to see all sides of the limb. Video interface boards developed by the Department of Radiology are used to determine the position of the laser line in 3D space. These boards and specially developed software allows shape acquisition on a low cost IBM/PC.Data from the scanner is then modified using an interactive 3D shape modification program.

These boards and specially developed software allows shape acquisition on a low cost IBM/PC.Data from the scanner is then modified using an interactive 3D shape modification program. This program called "Sockets" has been developed by Rehabilitation Medicine. This program allows the prosthetist to modify the computer generated limb shape in much the same way it is done manually. The limb shape can be moved in 3D space as it is being worked upon. There are functions to modify the shape to that of a proper socket. In addition there are measurement functions to check circumferences, distances and volume changes. There are viewing functions to allow the user to see the before and after shapes overlaid in both cross section and 3D wireframe. The shape can be viewed in both wireframe and solid shaded modes.

There are viewing functions to allow the user to see the before and after shapes overlaid in both cross section and 3D wireframe. The shape can be viewed in both wireframe and solid shaded modes. There are versions of "Sockets" for the Macintosh, IBM/PC, and Silicon Graphics computers.A plaster pattern for the socket is then created using a numerically controlled milling machine. The socket is produced by molding plastic over the pattern in the conventional manner.Using SLS to produce a socketThe prosthetic CAD data output from "Sockets" is not sufficient to produce a socket using selective laser sintering. Though geometric data comes from the prosthetic CAD program in the form of contours, the contours only describe the inside shape of the socket160and contain no information about attachments for fittings or the wall thickness of the socket.

In addition, the typical contour thickness is much higher than that required for SLS (0.125 inch versus 0.005 inch).A conversion program prQduces contours at the required resolution for the SLS system from prosthetic CAD data files. Software development was done on a Silicon Graphics 4D/25G. The conversion program allows input of wall thickness, a scaling factor, and the addition of an attachment fitting. The program outputs a data file in the contour file fonnat used by the SLS system. This contour file describes the polygonal outline of each layer of a part to be manufactured by the selective laser sintering process. The format is keyword driven and handles contours with multiple loops. The loops are oriented both clockwise and counter clockwise with respect to the z-axis. Outer loops are oriented counter clockwise, while inner loops, representing holes, are oriented in the clockwise direction.

Outer loops are oriented counter clockwise, while inner loops, representing holes, are oriented in the clockwise direction. If this convention is followed, the order of the loops within a contour definition is insignificant.The conversion program looks at the socket as being four separate parts. There are inner and outer socket walls and inner and outer fitting walls. The outer socket wall is created by scaling the inner wall appropriately to achieve the desired wall thickness. The fitting is made of inner and outer cylinders of fixed lengths. All the parts are represented by sets of vertical splines. Spline interpolation is used to produce contour slices of any desired thickness. The fitting is geometrically intersected with the bottom of the socket, producing the integrated socket and fitting.

Spline interpolation is used to produce contour slices of any desired thickness. The fitting is geometrically intersected with the bottom of the socket, producing the integrated socket and fitting. The outer socket wall and outer fitting wall are written as counter clockwise loops while the inner walls are written as clockwise loops.Comparison of Conventional Socket to Integrated Socket Produced with Selective Laser SinteringSocket with flattened bottom to accept fIXture ----...,....--.. FixtureFixture bolted to bottom of socket ---...... :r .....,integrated into bottom of socke~PylonConventionalSLS IntegratedFire2The integrated fitting is designed to accommodate a standard 30 mm pylon. A pylon location is completely reproducible as it is added during the design process rather than after socket construction. As can be seen in figure 2, the socket shape does not need to be distorted to accommodate a fitting.

As can be seen in figure 2, the socket shape does not need to be distorted to accommodate a fitting. The integrated SLS socket preserves the anatomical shape at the distal end of the limb.161ResultsThe SLS contour data files are sent by Internet from the Rehabilitation Engineering Lab at the Health Science Center at San Antonio to a Silicon Graphics workstation in the Mechanical Engineering Department at UT Austin. The contour files are converted to rasterized toggle point files. Several parts have been produced from prosthetic CAD data. These parts were scaled down so that they could be produced quickly. They were made using polycarbonate powder at a 5 mil thickness. All the parts so far have been made using the prototype machine on the UT campus.DiscussionMore sophisticated mounting fixtures will be developed in order to swiftly fit patients. These will be adjustable in order to change the alignment of the prosthesis.

These will be adjustable in order to change the alignment of the prosthesis. The mounting fixture is currently added by a postprocessor. The addition of the fixture will be incorporated into the prosthetic CAD program. The initial alignment will be done in the design phase rather than after socket construction. This is a very important step toward computer design of the entire prosthetic limb rather than just the socket.As experienced is gained new features will be tested. By varying the wall thickness, load bearing areas can be supported while retaining flexibility over sensitive areas of the residual limb. The cosmetic cover for the socket can be produced during socket fabrication. Other plastic compounds such as polypropylene will be tested as well.References1. AOPA Almanac: For Amputees ~ Walking to A Healthier Life. 37(1):22.2. Klasson B. Computer Aided Design Computer~Aided Manufacturer and other P & 0 IntI. 9:3~ 11, 1985.2. Klasson B.

brication of Smooth Surfaces based on Mask Projection Stereolithography Yayue Pan, Yong Chen*, Chi Zhou Epstein Department of Industrial and Systems Engineering University of Southern California, Los Angeles, CA 90089Corresponding author: yongchen@usc.edu, (213) 740-7829AbstractThe surface finish is critical for applications such as micro-fluid mixing and self-assembly that requires smooth fluidic flow and mechanical rotation respectively. However, in layer-based additive manufacturing processes, it is well known that the stair-stepping effect exists in the fabricated surfaces since a three-dimensional model is approximated by a set of two-dimensional layers. The fabricated surfaces are especially poor for the ones that are close to the horizontal plane. In this paper a novel approach for achieving improved surface finish has been presented for the mask-image-projection-based Stereolithography (MIP-SL) process.

In this paper a novel approach for achieving improved surface finish has been presented for the mask-image-projection-based Stereolithography (MIP-SL) process. Theoretical models and parameter characterization are presented with experimental verification. The developed approaches have been incorporated in the process planning of the MIP-SL process. Multiple test cases based on various types of curved surfaces have been performed. A comparison of the built results based on the traditional and the newly developed approaches has been discussed to illustrate the effectiveness of our method.Keywords: Additive manufacturing, surface finish, layer stair-stepping, mask projection stereolithography.1. IntroductionOne of the major problems in the layer-based additive manufacturing (AM) processes is the poor surface finish associated with the layer stair-stepping effect. As shown in Figure 1, a given three-dimensional (3D) model is first sliced into a set of two-dimensional (2D) layers.

As shown in Figure 1, a given three-dimensional (3D) model is first sliced into a set of two-dimensional (2D) layers. By stacking the 2D layers together, a physical part can be fabricated in an AM process to approximate the original computer-aided design (CAD) model. Due to the use of 2D layers, the fabricated part surfaces especially the ones whose normals are close to the building direction (Z axis) may have big approximation errors. Such poor surface quality limits the use of AM in applications that require smooth surfaces, e.g. the fabrication of 3D micro- channels in microfluidic systems or various assembly features in high-precision products.Figure 1. An illustration of the layer-based additive manufacturing processes and related stair-stepping effect.Since the approximation error depends on the layer thickness used in the slicing of a 3D model, a dominant approach for addressing the stair-stepping effect in AM is to reduce the layer thickness.

For example, the layer thickness typically used in a Stereolithography Apparatus (SLA) system is 0.1 mmwhile the layer thickness used in the latest developed inkjet-based systems (e.g. the ones from Objet Geometries Ltd) can be as small as 0.012mm. While the thinner layers in comparison always have less stair-stepping problems, the use of such ultra-thin layer thickness also significantly slows down the building process. In addition to the big sacrifice in building speed, it is also very difficult for AM processes such as SLA to significantly reduce its layer thickness since the spreading of liquid resin into uniform ultra-thin layers is well-know challenging.In this paper, we present an alternative approach for achieving smooth surfaces in the SLA- related processes. The AM process considered in this paper is the mask-image-projection-based stereolithograhy (MIP-SL). An illustration of the MIP-SL process is shown in Figure 1.a.

The AM process considered in this paper is the mask-image-projection-based stereolithograhy (MIP-SL). An illustration of the MIP-SL process is shown in Figure 1.a. The MIP-SL process is quite similar to the SLA process. However, a Digital Micromirror Device (DMD) instead of a laser is used in the MIP-SL process. A DMD is a microelectromechanical system (MEMS) device that enables one to simultaneously control ~1 million small mirrors to turn on or off a pixel at over 5 KHz. Hence mask images can be dynamically defined and projected on a surface area. By replacing a laser with a DMD, the MIP-SL process can be much faster than the SLA process.Figure 2. An illustration of the MIP-AM process and related approaches for improving its surface finish.Based on the Z axis and its surface normal N, all the surfaces in a 3D model can be classified into: (1) vertical surfaces (N • Z = 0), (2) down-facing surfaces (N • Z < 0), and (3) up-facing surfaces (N • Z > 0).

There is no need to specially consider the vertical surfaces since they generally do not contribute to the stair-stepping effect. For the other two types of surfaces, our approaches for improving their surface finishes are shown in Figure 2.b.(1) A gray scale image approach has been developed for the fabrication of smooth down-facing surfaces. As shown in Figure 2.b-(top), the key idea of the approach is to use a fine Z resolution to sample a given down-facing surface. Hence the curing depth at each pixel can be represented by a cure depth that has a higher resolution than the layer thickness. Accordingly a gray scale value at a pixel can be set to provide the desired energy input for the related curing depth. Consequently a gray scale image will be computed for all the down-facing surfaces in the MIP- AM process.(2) A meniscus equilibrium approach has been developed for the fabrication of smooth up-facing surfaces.

As shown in Figure 2.b-(bottom), the key idea of the approach is to closely match the fluid interfaces at the corners of intersecting planes to the related curved surfaces in the input geometry. Hence, liquid meniscus will be formed at the corners when the cured layers emerge outfrom the liquid. The shapes of the meniscuses depend on the contact horizontal and vertical surfaces, and can be modeled by considering the effects of capillarity, suction by gravity, multilayer adsorption and the boundary conditions [1, 2]. Accordingly a process optimization problem can be formulated in order to match the meniscuses to the given curved surfaces. Compared to the use of ultra-thin layers that significantly elongates the building time, the developed techniques enable a much bigger layer thickness to be used while achieving similar surface finish requirement. Hence significantly reduced building time could be expected in the MIP-AM process.

Hence significantly reduced building time could be expected in the MIP-AM process. To explain both techniques in more details, the remainder of the paper is organized as follows. The building technique developed for up-facing surfaces will be discussed first. Section 2 presents the models of formed meniscus in various cases. Accordingly the process planning for building smooth up-facing surfaces is presented in Section 3. The building technique developed for down-facing surfaces will then be presented. Section 4 presents the curing model for pixels with different gray scale values. Based on it, the process planning for building smooth down-facing surfaces is presented in Section 5. The experimental setup for performing physical experiments is presented in Section 6. The test results for various curved surfaces are discussed in Section 7. Finally, conclusion with future work is drawn in Section 8.2.

The test results for various curved surfaces are discussed in Section 7. Finally, conclusion with future work is drawn in Section 8.2. Modeling Formed Meniscuses in MIP-SLThe formation of the interfacial profile of the equilibrium meniscus that is attached to intersecting surfaces is first discussed. Such understanding and related mathematic models will provide the theoretical basis for the process planning and smoothness control of the up-facing surfaces. As shown in Figure 3.a, given a pair of walls, suppose the length of the horizontal plane is b, and the height of the vertical plane is h. With a finite horizontal plane, the maximum range of the fluid that can rise up on any x (=0,xlength b=)infinite vertical plane x y range of the fluid that can extend along an infinite horizontal plane (is hc. Similarly, with a finite vertical plane, the maximum 0,=ylength=h)is b0.yR2P1R1R1R2P2(a)(b)xzFigure 3.

Similarly, with a finite vertical plane, the maximum 0,=ylength=h)is b0.yR2P1R1R1R2P2(a)(b)xzFigure 3. Meniscus wetting to intersecting plane surfaces and fluid interface profile.Previous research on determining equilibrium shapes of fluid interfaces has been performed [3-7]. The Young-Laplace (YL) equation is the most widely accepted physical analysis of equilibrium meniscus shape. In the analysis, two types of forces, surface tension and pressure, are taken into account. The surface tension is defined with respect to a specific photopolymer resin. It is a thermodynamic quantity measuring the energy required to expand the area of the gas-liquid interface.

The surface tension is defined with respect to a specific photopolymer resin. It is a thermodynamic quantity measuring the energy required to expand the area of the gas-liquid interface. As shown in Figure 3.b, the YL equation relates the pressure difference ΔP across a fluid interface within a gravitational field to the curvature of the interface and the interfacial tension by:Δ =P P P 1 2−= −gy ρ γ 12=×(1 R 1+1 R 2); (2.1)where P1 and P2 are the pressures on either side of the interface between resin and air, g is gravity acceleration, y is the height of the meniscus above the horizontal plane surface, γ12 is the interfacial tension, R1 and R2 are the radii of curvature of the fluid interface. If R1 can be considered as infinite, the problem is simplified to a 2D case. Thus the following equation (2.2) can be derived from Eq.

If R1 can be considered as infinite, the problem is simplified to a 2D case. Thus the following equation (2.2) can be derived from Eq. (2.1):gy ρ γ− 12×(1+y && 2 2/3 y ) &=0;y &=dy dx;y &&dy = & dx(2.2)The interfacial tension γ has a relation with the capillary heightch and the contact angle θ asshown in Equation (2.3).ch2=2γ g ρ(1 sin ); θ−(2.3)So Equation (2.4) can be developed according to Equation (2.2) and (2.3).ρgy−1 2h c2×gy && (1 sin )(1 +ρ θ−y &2 2/3 )=0(2.4)2.1 Process Parameter CalibrationThe contact angle θ in Eq. (2.4) is the angle at which the liquid resin interface meets the solidified resin surface. Specific to the given liquid and solid system, the contact angle is determined by the interactions across the three interfaces, including liquid resin, solidified resin and air.

Specific to the given liquid and solid system, the contact angle is determined by the interactions across the three interfaces, including liquid resin, solidified resin and air. Researchers reported that θ is observed to be independent of volume and gravity and depends only on the surface tension [8, 9]. The capillary height hc is the maximum height that the fluid can reach on an infinite vertical wall. hc is a characteristic length for the fluid subject to gravity and surface tension.Both parameters (θ and hc) can be experimentally measured. Based on different b and h values, a set of experiments have been designed. Test parts with designed intersecting horizontal and vertical surfaces were built using the MIP-SL process. After curing the horizontal and vertical surfaces, the part is lifted up from the liquid resin slowly. Liquid meniscus in contact with the intersecting surfaces will be formed.

After curing the horizontal and vertical surfaces, the part is lifted up from the liquid resin slowly. Liquid meniscus in contact with the intersecting surfaces will be formed. After a long waiting time, the liquid volume will reach equilibrium over the horizontal wettable surface area. A mask image is then projected on the meniscus area to cure the liquid resin. The shape of the formed meniscuses can thus be captured in the built part. A set of parts with different sizes of horizontal and vertical planes have been built. Two of such built parts are shown in Figure 4. Based on the experiments, the contact angle θ and the capillary height hc can be measured. For the liquid resin used in our experiments (envisionTEC SI500, Ferndale, MI), we estimated that θ =25o, and hc =1.40mm.Figure 4.

For the liquid resin used in our experiments (envisionTEC SI500, Ferndale, MI), we estimated that θ =25o, and hc =1.40mm.Figure 4. Two built parts with different b and h values for measuring surface tension parameters.2.2 Meniscus Shape Analysis and SimulationIn the MIP-SL process, let h denote the height of the vertical plane and b be the length of the horizontal plane. hc and b0 denote the maximum range the liquid can reach on the vertical and horizontal plane respectively. The formed meniscus shapes for different b and h values can be analyzed based on the following scenarios: h>;c h b b 0 ;c h b b > ; 0 ;c h b b 0 ;c h b b < . 0>(1) Case 1:;h<(2) Case 2:h>< ; and(3) Case 3:h<(4) Case 4:Different boundary conditions can be derived for each scenario, which are discussed as follows.(1) Case 1: both h and b can be considered as infinite.

The related boundary conditions are: o tan 25( y x &=0)= −ctancθ = −= −2.14129;(0)xb= = ∞ (2.5)Using the governing equation (2.4) and boundary conditions (2.5), the height of the fluid can reach on the vertical wall is calculated to be 1.398372mm, which agrees well with the experiment result of 1.40 mm. We used Matlab (MathWorks, Natick, MA) to solve this problem and plotted the result as shown in Figure 5.a.mmmmmmmm(0.5 < h < 1.4)mmmm(a) Case 1(b) Case 2(0.1 < h < 0.5)Figure 5. Plotting results of the meniscus profile in cases 1 and 2.(2) Case 2: h is smaller than hc(x= ∞ ), and b can be considered as infinite. Hence the curvature of the meniscus is decided by h.

Plotting results of the meniscus profile in cases 1 and 2.(2) Case 2: h is smaller than hc(x= ∞ ), and b can be considered as infinite. Hence the curvature of the meniscus is decided by h. The boundary conditions are as follows.( y x & ( y y &==0) 0)= −= −c ctan tancθ = − cθ ≤ −o tan 25 o tan 65= −= −2.14129; 0.466038;(0)yh= ;(2.6)In this case, the length of the fluid extended on the horizontal plane b0 is dependent on the value of h, namely, b0= f(h). The meniscus shapes with different h were estimated by solving the governing equation (2.4) and the boundary conditions described in equation (2.6) using Matlab. The plotting results of different h values are shown in Figure 5.b.(3) Case 3: h is bigger than hc(x= ∞ ) and b is smaller than b0. Thus the curvature of the meniscus is decided by b.

The plotting results of different h values are shown in Figure 5.b.(3) Case 3: h is bigger than hc(x= ∞ ) and b is smaller than b0. Thus the curvature of the meniscus is decided by b. The boundary conditions are as follows.y x ( & y y ( = & y b ( ) 0 ==0)0)= −= −cctantanθθ= −≤ −cctan 25 otan 65 o= −= −2.14129;0.466038;(2.7)Based on the governing equation (2.4) and the boundary conditions Equation (2.7), the meniscusshapes can be computed in Matlab. Figure 6 shows the plotting results for different b values.mmy=f(x); x(0) = 5; x(0) = 2; x(0) = 0.9; x(0) = 0.5mmFigure 6. Plotting result of meniscus profile in Case 3.(4) Case 4: h is smaller than hc(x= ∞ ), and the length of the horizontal plane b is smaller than b0. Hence the curvature of the meniscus is decided by both b and h. The problem follows the boundary conditions:y = (2.8) The meniscus shape can be calculated for different b and h values.

Hence the curvature of the meniscus is decided by both b and h. The problem follows the boundary conditions:y = (2.8) The meniscus shape can be calculated for different b and h values. Three examples by using(0)=, h y b( ) 0Matlab to solve the mathematic model and the related boundary conditions are shown in Figure 7.mmmmmmmmmmmmFigure 7. Plotting result meniscus profile in Case 4.3. Process Planning for Building Smooth Up-facing SurfacesSuppose an input 3D CAD model with curved up-facing surfaces is given as shown in Figure 8.a. The traditional additive manufacturing method will slice the model into multiple 2D layers based on a given layer thickness. One mask image will be generated for each layer in the MIP-SL process to cure the related 2D shape. Consequently the built part will inevitably have stair-stepping defects.(a)(b)Figure 8.

One mask image will be generated for each layer in the MIP-SL process to cure the related 2D shape. Consequently the built part will inevitably have stair-stepping defects.(a)(b)Figure 8. Comparison between the traditional and the meniscus equilibrium methods.In comparison, Figure 8.b illustrates the major steps of the meniscus equilibrium method for achieving smoother up-facing surfaces. For a specific layer i, c(x, y) is computed which represents the ratio of the Z-height at the location (x, y) to the layer thickness LT. Similar to [10], a sliced image with gray scale values can be generated for each layer. Accordingly the thickness at pixel (x, y) is c(x, y)×LT. Hence instead of a binary layer image, the sliced image has pixels with ratios c(x, y) which are continuous and c(x, y) ∈ [0, 1].

Accordingly the thickness at pixel (x, y) is c(x, y)×LT. Hence instead of a binary layer image, the sliced image has pixels with ratios c(x, y) which are continuous and c(x, y) ∈ [0, 1]. The original gray scale image can be split into two mask images such that the mask image1 can cure a portion of layer i that will be used in forming meniscus while the mask image2 can be used in curing the formed meniscus. Note it is not necessary to use the meniscus equilibrium approach for each layer. Depending on the given geometry, meniscuses can be formed and cured after multiple layers. As discussed in the meniscus mathematical model, two important parameters in determining the shape of the meniscus profile are h and b. Thus the planning of image1 and image2 can change the setting of h and b, which lead to different approximation errors between the meniscus profiles and the desired shapes. An illustration example is given in Figure 9.a.

An illustration example is given in Figure 9.a. For given shape boundary as shown in black lines, two sets of mask images (image1 and image2) can be used, which define two vertical walls as shown in brown lines. Accordingly the formed meniscuses will have different profiles due to different b or h values as shown in Figure 9.a.In order to generate desired MIP-SL process plan to ensure the meniscus shapes match the given curved geometry, the process planning for the up-facing surfaces can be mathematically defined. Two types of errors need to be considered including the shape and curvature approximation errors. The shape approximation errors εa can be computed by comparing the difference between the meniscus profile f(x)and the desired geometry profile F(x). The curvature approximation error εc can be computed by comparing the first derivative values of f(x) and F(x).

y ∈)=)F x ( 2−(;b F x i 1=−x 2x 1h i(3.1)(0, ),h x∈(x x , 1 2)) is given. The input curve can be divided into several pieces such that h in the process planning is smaller than the maximum vertical height hc as discussed in the meniscus model. Suppose LT is the layer thickness and a sampling distance e is used in computing the errors of the meniscus profile. An algorithm based on the greedy heuristic can be developed as follows. The meniscus shape in the area of is first hP (i.e. estimated and compared with the input geometry. If the errors are within the acceptable ranges, point 3) can be selected as the meniscus point; otherwise, the meniscus shape in the area of 0hP is found that satisfies the given y approximation error ranges (e.g. point 1). The remaining curve can then be processed in a similar way to hiP (e.g. points 1, 2, 3).

Modeling Curing Depths Related to Gray Scale Pixels in MIP-AMThe meniscus equilibrium method can only address the up-facing surfaces. To improve the surface finish of down-facing surfaces, a gray scale mask image method has been developed, which is motivated by the idea of the optimized mask image projection [11]. As extensively studies in the SLA process [12], the critical threshold (Ec) is a primary parameter for a given photosensitive resin. The liquid resin will be solidified when the exposed energy is bigger than Ec; otherwise, the resin will remain as liquid or gel. In addition, the curing depth (Cd) is directly related to the energy input E if E > Ec. The dependence of Cd as a function of the exposure E is generally log-linear [12]. Their relation follows the equation:Cd=Dpln(E E c)(4.1)One way to control E is to adjust the exposure time for a binary projection image [13].

Their relation follows the equation:Cd=Dpln(E E c)(4.1)One way to control E is to adjust the exposure time for a binary projection image [13]. Another approach is to adjust pixels’ gray scale values while fixing the exposure time of the projection image. In comparison, the gray scale image approach is easier to control. Hence, in the MIP-AM process, full light intensity will be projected for a pixel of a mask image if its gray scale value g = 255; when 0 < g < 255, only partial light intensity will be projected for curing the resin area related to the pixel. Corresponding to a given Ec and an exposure time T, a threshold gray scale value gc can be identified. An illustration of the gray scale image method is shown in Figure 10. Suppose gc = 75. Hence the exposures at pixels B, C and D whose gray scale values are set as 102, 153 and 204 respectively will lead to cured shapes with different cure depth.

Suppose gc = 75. Hence the exposures at pixels B, C and D whose gray scale values are set as 102, 153 and 204 respectively will lead to cured shapes with different cure depth. No resin will be cured at pixel A due to its insufficient energy input. In addition, the cure depth of a pixel is required to be bigger than a layer thickness such than sufficient over-cure exists. Hence the current layer can be bonded to the previous layers.Figure 10. Principle of the gray scale image method - the cure depth varies with different gray scale values.For the liquid resin and the light source used in our testbed, designed experiments have been performed to calibrate the relations between the cure depth and the gray scale values. The CAD model of a designed part is shown in Figure 11.a. In the test, the exposure time was set at 3 seconds. Mask images with different gray scale values were used in curing the top layer.

The CAD model of a designed part is shown in Figure 11.a. In the test, the exposure time was set at 3 seconds. Mask images with different gray scale values were used in curing the top layer. The part was then taken out and the related cured layer thickness Cd was measured. The recorded Cd values for different gray scale values are shown in Figure 11.a. The cure depth increases with the set gray scale value. In addition, the gray scale value threshold gc is around 76 when T = 3 seconds. For a pixel with g < gc, liquid resin was not cured or the cure depth was too small for a given layer thickness (i.e. < 0.1mm). The relation between Cd and ln(g) has also been computed as shown in Figure 8(c).

< 0.1mm). The relation between Cd and ln(g) has also been computed as shown in Figure 8(c). The relation matches Equation 4.1 quite well and can be approximated by the following equation:F x ( )⎧ = ⎨ ⎩g 0.521*ln( ) 2.252,−0,g<gcg≥gc;(4.2)Top layerA)m m ( h t p e D e r u CA0.70.60.50.40.30.20.1Cure Depth)m m(h t p e D e r u C0.60.50.40.30.20.1Physical Cure Depth Fitted Cure Depth0255239223207191 143 Gray scale values Gray Scale Value1751591281129604.55Ln (Gray Scale Value g)5.5(a) Experiments on cd and g values(b) Relation of cd and ln(g) valuesFigure 11. Relation between the cure depth with the gray scale value for SI 500 resin with the exposure time of 3s.Hence by slightly changing the gray scale values of neighboring pixels in a projected mask image, the cure depth related to the pixels can be gradually changed.

Thus a higher Z resolution of down-facing surfaces can be achieved, which will lead to improved surface finish. In the process planning for a given 3D model with down-facing surfaces, the cure depth at each pixel needs to be accurately controlled by adjusting its gray scale value. In addition, an upward Z offsetting distance needs to be considered since the range of achievable cure depths (i.e. 0.1mm < d < 0.65mm) is bigger than a layer thickness to ensure the over-cure between layers. Such Z offsetting of down-facing surfaces has also been used in the SLA process [12].Figure 12. Comparison between the traditional and the gray scale image method.5. Process Planning and Smoothness Control for Down-facing SurfacesThe schematic illustration of the traditional and the gray scale image methods are shown in Figure 12. Several layers are reserved as compensation for over-cure. In the traditional method, a set of layers are sliced based on the Z offsetting result.

Several layers are reserved as compensation for over-cure. In the traditional method, a set of layers are sliced based on the Z offsetting result. Accordingly binary mask images are used in the MIP-AM process. In the gray scale image method, a compensation Z thickness is calculated based on the minimum and maximum curing depth for gray scale pixels (i.e. 0.1mm and 0.65mm respectively in our test). Hence the desired thickness in each layer is bigger than the minimum achievable curing depth and smaller than the maximum achievable curing depth. Based on the Z offsetting result and the established relation between curing depth and the gray scale values, a mask image with desired gray scale values at all the pixels can be computed. The mask image can then be used in the MIP-AM process for building smooth down-facing surfaces. Figure 13.a shows the related algorithm for generating gray scale mask images.

The mask image can then be used in the MIP-AM process for building smooth down-facing surfaces. Figure 13.a shows the related algorithm for generating gray scale mask images. An example is shown in Figure 13.b to illustrate the two main steps, Z offsetting and gray scale image setting.ZLayer ThicknessMax CdZ offset boundaryΔdSlicing planesPortion to be curedCd1Cd3Gray scale mask imageg1g2g3g4g5(a) Main steps in the gray scale image method(b) An illustration exampleFigure 13. A gray scale image generation algorithm with an illustration example.6. Experimental SetupA prototype system has been built for verifying the presented approaches. The hardware setup of the developed MIP-SL system is shown in Figure 14.a. In the designed system, an off-the-shelf projector (Acer H5360) was used as the projection device. The use of a commercial projector can significantly reduce the prototype cost and simplify the system design.

The use of a commercial projector can significantly reduce the prototype cost and simplify the system design. The optical lenses of the projector were modified to reduce the projection distance. Various projection settings including focus, key stone rectification, brightness and contrast were adjusted to achieve a sharp projection image on the designed projection plane. The DMD resolution in our system is 1024×768 and the envelope size is set at 110×82 mm. Two linear motion stages from Thomson Industries Inc (Radford, VA) are used in the system. One is used as the Z elevator for lifting the build platform; another is used as the sweeper for moving a blade to flatten the resin surface. A high performance 8-axis motion control board KFLOP+SnapAmp 1000 (Dynomotion Inc., Calabasas, CA) was used for driving the linear stages.Input boundaryXCommercially available photocurable resin (envisonTEC SI500, Ferndale, MI) was used in the experiments.

The exposure time was set at 3 second based on the curing depth analysis. A mask imaging planning and control software system has been developed by using the C++ programming language with Microsoft Visual C++ compiler. The graphical user interface (GUI) of the developed software system is shown in Figure 14.b. The system can load in a STL model, perform process planning and synchronize the mask image projection with desired linear stage motions.Figure 14. Hardware and software setup of the developed MIP-SL testbed.7. Results and DiscussionsA set of test cases have been designed to test the meniscus equilibrium method for up-facing surfaces and the gray scale method for down-facing surfaces. The experimental results are presented as follows.7.1 Tests on Straight Up-facing SurfacesAs shown in Figure 15.a, a CAD model with a slanted surface M was used in this test.

The experimental results are presented as follows.7.1 Tests on Straight Up-facing SurfacesAs shown in Figure 15.a, a CAD model with a slanted surface M was used in this test. To compare the differences between the traditional and newly developed methods, the surface M was split into two halves, M1 and M2. The traditional method was used in building the upper half M1, and the meniscus equilibrium method was used in building the bottom half M2. The same layer thickness (0.1mm) was used in the test. Accordingly a set of mask images were generated by slicing the given CAD model. However, one meniscus point every ten layers was added in set {M} during building M2. That is, after every 10 layers have been built, the built part will be raised up by a certain distance (e.g. 1.5mm). After waiting for several minutes, liquid will reach equilibrium. A mask image2 related to the portion of M2 within the built 10 layers is then projected to cure the formed meniscuses.

1.5mm). After waiting for several minutes, liquid will reach equilibrium. A mask image2 related to the portion of M2 within the built 10 layers is then projected to cure the formed meniscuses. The building result is shown in Figure 15.a. It is obvious that the appearance of M2 portion is much smoother than that of the M1 portion.Quantitative measurements have been performed to better understand the surface quality difference. The Z height of a set of uniformly sampled points on M was measured by using a digital height gauge with a 0.1mm probe tip. To automate the measuring process, two motion controlled linear stages (Velmex Inc, Bloomfield, NY) were used. In the measurement setup, the built part was fixed on a linear stage to be translated in the X axis. The digital height gauge was fixed on another linear stage to be translated in the Z direction. Every time when the probe touches a sampling point on the surface, a height value can be recorded from the digital gauge.

Every time when the probe touches a sampling point on the surface, a height value can be recorded from the digital gauge. To avoid the friction between the probe tip and the slanted surface, the gauge is move away from M in Z direction first before the part is fed by a small distance in X direction. The recorded Z height values of the sampling points on surface M is plotted in Figure 15.b. The X axis denotes a set of uniform-spaced sampling points, where 0 denotes the starting point of the measurement and 400 denotes the 400th measuring point. The Z axis is the readouts from the digital height gauge. The recorded data of the upper half M1 and the bottom half M2 are plotted by using blue and redpoints respectively. The quantitative comparison shows that the stair-stepping effect of M1 is much bigger than that of M2. We also marked three portions, a1, a2 and a3, on the surface as shown in Figure 15.b. They represent M1, the transition between M1 and M2, and M2 respectively.

We also marked three portions, a1, a2 and a3, on the surface as shown in Figure 15.b. They represent M1, the transition between M1 and M2, and M2 respectively. A precision measurement machine (Sol system from Micro Vu Inc., Windsor, CA) was used in taking microscopic images of these three portions. The captured images are shown in Figure 15.c. The results illustrate that the traditional method will lead to a ragged surface, while the meniscus equilibrium method can effectively eliminate the stair-stepping effect in the MIP-SL process.CAD ModelMTop ViewM1 M2Z (mm)Side ViewM1M2Sampling points(a) CAD model and built part(b) Surface quality measurement resultsa1a2a30.5 mm0.5 mm0.5 mm24(c) Microscopic images of three different portions of the built up-facing surface Figure 15.

A comparison of the built straight up-facing surface based on different methods.7.2 Tests on Curved Up-facing Surfaces – Concave CasesAs shown in Figure 16.a, a CAD model with a curved surface M was used in this test. The same process parameters were used in the traditional and newly developed methods for building the CAD model. The related built results M1 and M2 are shown in Figure 16.a. The same measurement procedure as described in Section 7.1 was followed for a quantitative comparison of their surface finishes. The Z height values of the curved up-facing surface are shown in Figure 16.b. The microscopic images of a portion of M1 and M2 are shown in Figure 16.c.The meniscuses in this test case wet to a cylinder and an intersecting planar surface. The test results verify the capability of the meniscus equilibrium method in fabricating smooth concave up-facing surfaces.

The test results verify the capability of the meniscus equilibrium method in fabricating smooth concave up-facing surfaces. One limitation of the developed method is that the given geometry should have an outlet for the liquid resin to flow away such that meniscuses can be formed. In the process planning of the meniscus equilibrium method, there may also be discontinuous error between neighboring meniscuses as shown in Figure 16.a. In order to reduce such error, the generated masks image1 and the layers at which image2 will be projected can to be adjusted. However, reducing the discontinuous errors may increase the shape approximation error. Hence a balance between the two types of errors needs to be considered. Computing process plans to reduce both errors will be further studied in our future work.CAD ModelTop ViewM176.5M16M25.55M24.5Side ViewM243.532.5Discontinuous Discontinuous error error20100200300400500M10.5 mmFigure 16.

A comparison of the built concave up-facing surfaces based on different methods.7.3 Tests on Curved Up-facing Surfaces – Convex CasesAs shown in Figure 17.a, a CAD model with a curved surface M was used in this test. The same process parameters were used in building the model by using the traditional and newly developed methods. The related build results M1 and M2 are shown in Figure 17.a. The same measurement procedure as described in Section 7.1 was followed for a quantitative comparison of their surface finishes. The Z height values of the curved up-facing surface are shown in Figure 17.b. The microscopic images of a portion of M1 and M2 are shown in Figure 17.c. The test results verify the capability of the meniscus equilibrium method in fabricating smooth convex up-facing surfaces.7.4 Tests on Down-facing SurfacesAs shown in Figure 18.a, a CAD model with a down-facing surface was used in this test.

The same process parameters were used in building the model by using the traditional and newly developed methods. To generate gray scale images for building the part, the CAD model was offset by 0.3mm. Accordingly the gray scale value at each pixel was calculated based on the corresponding curing depth. Consequently the computed gray scale images can cure a layer of liquid resin with uneven thicknesses. The related build results M1 and M2 are shown in Figure 18.a. The same measurement procedure as described in Section 7.1 was followed for a quantitative comparison of their surface finishes. The Z height values of the down-facing surface are shown in Figure 18.b. The microscopic images of a portion600M20.5 mm26of M1 and M2 are shown in Figure 18.c. It is obvious from the images that the traditional method can only cure layers with an increament of the given layer thickness. In comparison, the gray scale image method can achieve varying Z heights within one layer.

In comparison, the gray scale image method can achieve varying Z heights within one layer. Therefore, the gray scale image method can significantly improve the surface smoothness of down-facing surfaces in the MIP-SL process.Top ViewM1M2Z (mm)9.59M18.5M287.576.56CAD ModelSide View5.55050100150200250300 Sampling points350(a) CAD model and built part(b) Surface quality measurement resultsM2M1m m 5 0.m m 5 . 0Figure 17. A comparison of the built convex up-facing surfaces based on different methods.7.5Z (mm)Top ViewM17.37.1M1M2CAD Model6.96.76.56.36.15.95.7Sampling points5.5050100150200250Side ViewM2(b) Surface quality measurement resultsDown-facing surfacesDown-facing surfacesM1M2(a) CAD model and built parts(c) Microscopic images of the built surfacesFigure 18. A comparison of the built down-facing surfaces based on different methods.8.

A comparison of the built down-facing surfaces based on different methods.8. ConclusionsThis paper presented a unique approach to build smooth curved surfaces in the mask-image-projection- based stereolithography process. A meniscus equilibrium method and a grey scale image method have been developed for up-facing and down-facing surfaces respectively. Based on the developed meniscus models in various scenarios, a process planning problem has been formulated for up-facing surfaces. Accordingly an algorithm based on the greedy heuristic was developed for the process planning problem. Based on the calibrated model between curing depths and gray scale values, a process planning algorithm has been developed for down-facing surfaces. Experimental tests based on the developed methods have been performed and compared with the traditional method.

Experimental tests based on the developed methods have been performed and compared with the traditional method. The differences in the build results illustrate the effectiveness of our method in improving the surface finish of curved surfaces in the MIP-SL process. Consequently it is possible to build smooth surfaces without sacrificing building time.The presented method could be further improved to minimize the shape and curvature approximation errors in an up-facing surface. Such improvements can be made by developing better solution strategies for the formulated process planning problem.References1. Elena, D. M., Fuentes, J., Cerro, R. L., and Savage, M. D. (2010), "An analytical solution for a partially wetting puddle and the location of the static contact angle", Journal of colloid and interface science, Vol. 348, no. 1, pp. 232-239.2. Neumann, A.W., Benhabib, B., Szekely, J., Racz, L. M., and Rooks, S.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFABRICATION OF SUPPORT-LESS ENGINEERED LATTICE STRUCTURES VIA JETTING OF MOLTEN ALUMINUM DROPLETSDinesh Krishna Kumar Jayabal*, Khushbu Zope*, and Denis Cormier*Department of Industrial and Systems Engineering, Rochester Institute of Technology, Rochester, NY 14623AbstractMagneto Hydro Dynamic (MHD) jetting is a promising new metal additive manufacturing technique that employs on-demand jetting of molten metal droplets onto a moving substrate. A particularly unique aspect of the process is its potential to print down-facing features without the need for support structures. Under suitable droplet jetting conditions affecting time and temperature, each droplet at least partially solidifies prior to impact of the next incoming molten metal droplet.

Under suitable droplet jetting conditions affecting time and temperature, each droplet at least partially solidifies prior to impact of the next incoming molten metal droplet. The combination of droplet jetting frequency and substrate velocity dictates the step- over distance between incoming droplets. With relatively large droplet step-over distances (or equivalently small percentage of droplet overlap), it is possible to print unsupported down-facing features that are nearly parallel to the X-Y build platform. In this paper, we describe initial results in which engineered lattice structures have been printed using 4043 aluminum using this approach. A parametric study that maps jetting frequency and droplet step-over distance with the resulting lattice strut angle is presented.

A parametric study that maps jetting frequency and droplet step-over distance with the resulting lattice strut angle is presented. With careful control of jetting parameters, we show that it is possible to print nearly horizontal lines without any support.IntroductionThe majority of metal additive manufacturing research to date has focused primarily on powder bed fusion processes (e.g. DMLS, SLM, EBM, etc.), binder jetting of metals, or directed energy deposition. In all three of these process categories, the feedstock material used is typically metal powder. Powder-based processes are extremely versatile in the sense that most metals can readily be atomized into powder form. However, metal powders are typically much more expensive than the same alloy in wire or bar form, they require considerably more safety precautions than metal wire, and the high surface area-to-volume ratio makes them very susceptible to oxidation.

Conversely, not all alloys can be easily drawn into wire form. The range of alloys available in wire form is therefore not as great as the range of alloys available in powdered form. Nevertheless, there has been a considerable growth of interest in wire-based metal additive manufacturing processes.Examples of wire-based metal additive manufacturing include electron beam freeform fabrication (EBF3) [1] and wire and arc additive manufacturing (WAAM) [2]. These processes feed wire into a melt puddle created by an electron beam or arc welding head. A new droplet jetting process that uses wire feedstock material is magneto hydro dynamic (MHD) liquid metal droplet jetting [3]. The MHD process melts wire in a micro crucible. The molten metal is directed into a pump chamber surrounded by a coil that induces eddy currents in the molten metal.

The MHD process melts wire in a micro crucible. The molten metal is directed into a pump chamber surrounded by a coil that induces eddy currents in the molten metal. Lorenz forces induce jetting of droplets at a specified frequency from the nozzle.757This paper is focused on the use of MHD droplet jetting to produce engineered lattice structures. With traditional powder bed fusion techniques, downward-facing surfaces of lattice struts are supported by unfused powder which is then removed after part fabrication. This generally works just fine if the lattice unit cell size is sufficiently large. Powder removal from lattice structures with small unit cell sizes can be a challenge though. Likewise, there is typically a desire to have components with completely solid skins. In powder bed fusion, a solid-skinned part with an engineered lattice core cannot be produced because there would be no way to remove unfused powder.

In powder bed fusion, a solid-skinned part with an engineered lattice core cannot be produced because there would be no way to remove unfused powder. In this paper, initial results are presented on research in which the MHD liquid metal jetting process developed by Vader Systems is used to fabricate engineered lattice structures without any support structures.Experimental SetupFor this study, a Vader Systems MK1 drop-on-demand metal AM machine was used. The MK1 is based on the aforementioned magnetohydrodynamic jetting process. The printer setup consists of a spool of solid metal wire (4043 Al in this case) which is fed into a resistively heated chamber where it melts. A coil for electromagnetic pulsing results in Lorenz forces that cause ejection of discrete droplets from a ceramic nozzle. A level sensing laser system maintains a constant level of molten metal in the reservoir, and a thermocouple provides temperature control.

A level sensing laser system maintains a constant level of molten metal in the reservoir, and a thermocouple provides temperature control. As metals form oxides at elevated temperatures, a constant stream of argon is used to form an inert gas sheath around the jet. The components of the print head are shown in Figure 1.Figure 1: Vader MK1 print head componentsWith the MHD process, molten metal droplets are ejected through the nozzle by creating a pressure pulse within the molten metal pool inside the reservoir with help of a pulsed electromagnetic field. The pulsed field induces a transient current in the pool and as this induced current is under the applied field, the resulting Lorentz force causes the ejection of molten metal758droplets from the reservoir through the nozzle orifice. The characteristics of the pulsed field in question is dependent upon the input voltage and the waveform produced by the function generator.

The characteristics of the pulsed field in question is dependent upon the input voltage and the waveform produced by the function generator. The frequency of the droplets ejected is controlled by the pulsing frequency of the coil. In the Vader Systems MK1 machine, the print head assembly moves up/down in the Z-axis direction, while the print bed moves in the X-Y plane. The timing of droplet ejection is coordinated with motion in the X, Y and/or Z-axis directions in order to build the desired 3D object. The process parameters that were held constant for the experiments in this research are provided in Table 1.Table 1: Process parameters for the experiment.Parameter Reservoir temperature (°C) Print bed temperature (°C) Argon flow rate (SCFH) Nozzle diameter (µm)Value 950 350 15 500By design, the Vader MK1 machine generates drops only when there is a movement in the Cartesian space.

By choosing a value for drop spacing, one can instruct the machine to generate a drop every time it moves a distance matching the value set for the drop spacing. Here, the drop spacing is expressed through a drop overlap fraction for a given drop size as shown in the equation below.𝐷𝑟𝑜𝑝 𝑠𝑝𝑎𝑐𝑖𝑛𝑔 = 𝐷𝑟𝑜𝑝 𝑑𝑖𝑎𝑚𝑒𝑡𝑒𝑟 × (1 − 𝐷𝑟𝑜𝑝 𝑜𝑣𝑒𝑟𝑙𝑎𝑝 𝑓𝑟𝑎𝑐𝑡𝑖𝑜𝑛)……………… Equation 1Dropwise solidification occurs when droplets are deposited at a frequency low enough for one droplet to solidify before the next droplet arrives. If a newly arriving droplet’s position is offset some amount relative to the last droplet, then an inclined column will form as droplets stack up. The degree of droplet overlap will affect the angle of the printed column. Likewise, the droplet jetting frequency will affect droplet solidification and spreading behavior. At higher frequencies, the columns tend to get shorter as a result of higher heat content and a greater degree of droplet spreading [4].

At higher frequencies, the columns tend to get shorter as a result of higher heat content and a greater degree of droplet spreading [4]. It is also to be noted that even though the nozzle diameter used in this research was 500 µm, the actual droplet size tends to be slightly smaller than this. As the strategy for building lattice structures is entirely dependent on the drop size, it is essential to measure drop size as a function of jetting parameters.Parameter MappingFor the droplet deposition experiment to print slanted pillars, factors such as droplet temperature, substrate temperature, droplet jetting frequency, stand-off distance between the print head and the substrate, and the droplet overlap fraction (which is the percentage fraction of the overlap between the two successive droplets (Equation 1) were considered. To understand the effect of these factors on the angles of the printed pillars, a full factorial study was performed.

To understand the effect of these factors on the angles of the printed pillars, a full factorial study was performed. The factors and levels considered for this study are listed in Table 2.Pillars printed at different process parameter settings are depicted in Figure 2. Each of the nine images depicts a set of pillars printed with droplet overlap fractions ranging from 0.5 to 1.759With a droplet overlap fraction of 1 (i.e. droplets overlap completely), a vertical pillar is formed at a 90° angle with respect to the substrate. As the overlap fraction decreases (i.e.

droplets overlap completely), a vertical pillar is formed at a 90° angle with respect to the substrate. As the overlap fraction decreases (i.e. there is less overlap between droplets), the angle of the pillar with respect to the substrate decreases.Table 2: List of factors and levels used for the parameters mapping experimentFactor Stand-off distance (mm) Frequency (Hz) Droplet overlap fractionLevels45, 65, and 85 9, 12, 15, 18, 21, 24, 27, and 30 0.5 to 1 in steps of 0.025 (21 levels in total)Figure 2: Printed pillars under different process parameters760As droplet jetting frequency was increased, the temperature increases, and droplet spreading increases. This generally results in shorter pillars with a larger diameter. However, with increased frequency, smaller droplet-overlap-fractions did not yield the linear slanted pillars but yielded wavy non-uniform pillars (Figure 3). At higher nozzle stand-off distance, more deformities were observed in the slanted pillars.

At higher nozzle stand-off distance, more deformities were observed in the slanted pillars. The slanted pillars printed at every combination of frequency and stand-off distance were captured in images. These images were used to determine the angles between the slanted pillars and the substrate using ImageJ [5]. against Collected data was plotted frequency and droplet overlap fraction (Figure 4). Compared to stand-off distances of 45 and 85 mm, a 65 mm stand-off distance yielded more standing pillars. This can be explained on the basis of conditions. Good forming optimum pillar mechanical bonding between the two successive droplets can be achieved if the incoming droplet carries the sufficient thermal energy to re-melt the top surface of the previously deposited droplet.

If this incoming thermal energy per unit time with the droplets is too low, as could happen in the case of low jetting frequency or/and higher stand-off distance, then either inter-droplet bonding will be weak or there will be no bonding at all. On the other hand, if the incoming thermal energy per unit time with the droplets is too high (e.g., high frequency or/and then one deposited droplet has insufficient time to begin solidifying before the next droplet lands. In this case, the pillar slumps or does not form at all. Thus, to produce linear slanted pillars at a desired angle for a given engineered lattice structure, the droplets carrying the correct amount of thermal energy will result in an appropriate amount of remelting at the overlapping top surface of previously deposited droplets. As droplet overlap fraction increases, the angle formed with the substrate increases.

As droplet overlap fraction increases, the angle formed with the substrate increases. A large set of parameter combinations at 65 mm stand-off distance resulted in suitable thermal energy for in subsequent thus was used bonding and experiments to print complete lattice structures.lower stand-off distance),761Figure 3: Correlation map between the frequency, droplet overlap fraction, and the angle formed by pillar with the substrate at three different stand-off distances.Printing Lattice StructuresFigure 4: Cuboid geometries of the lattice unit cells with square shaped base area (w x w) and cell height (h), where α is the angle made by the slanted pillar with the bottom surface of the cuboid and α1 < α2 < α3.the knowledge obtained from the parameters mapping three different cuboid experiment, lattice unit cells were selected for the printing experiment (Figure 4 and lattices Table 3). The printed corresponding to parameter listed in Table 3 are shown in Figure 5.

The printed corresponding to parameter listed in Table 3 are shown in Figure 5. Angles mentioned in Table 3, were calculated from the dimensions of the lattice structures depicted in Figure 5.Based onTable 3: Printing parameters and geometrical dimensions of the lattice structuresDimension or ParameterUnit Cell Width, w (mm) Unit Cell Height, h (mm) Diagonal Strut Angle, α (°) Frequency (Hz) Stand-off distance (mm) Droplet overlap fraction Droplet mass (µg) Droplet diameter (µm)Unit Cell 1 5.3 2.9 21.15 12 65 0.65 105 420Unit Cell 2 6.1 6.3 36.14 21 65 0.8 148 472Unit Cell 3 6 8.7 45.71 27 65 0.8 126 447Figure 5: Lattice structures (3X3X2) with unit cells 1, 2, and 3. The printing parameters for these lattice structures were same as that for unit cells 1, 2, and 3, respectively.762A program was written to generate the G-code for printing lattice structures having the desired configurations and cell geometries.

A subroutine in the program was written to print inclined pillars in the coordinates dictated by the unit cell geometry and the number of unit cells in the lattice structure. The generated G-code file is automatically converted to a file with an .mpf extension, the type which Vader MK1 uses for printing parts.Stand-off distance was held constant at 65 mm for all three lattice prints. The droplet overlap fraction used in the unit cell 1 sample was smaller than that used for the unit cell 2 and 3 samples in order to produce a shallower strut angle. Likewise, the unit cell 1 sample was printed with a lower drop jetting frequency in order to reduce sagging, or drooping, of pillars due to the more shallow strut angle. The unit cell 2 and 3 samples were printed at same droplet overlap fraction with slight modification in the frequency. This was possible because of the different droplet masses.

The unit cell 2 and 3 samples were printed at same droplet overlap fraction with slight modification in the frequency. This was possible because of the different droplet masses. The droplet mass, and thus the droplet diameter, for the unit cell 2 sample was higher than that for the unit cell 3 sample. Larger diameter droplets carry higher thermal energy towards the pillar being printed and spread more than a comparatively smaller and lighter droplets. Thus, bigger droplets will facilitate more spreading of the droplets in the same linear length of the pillar. In the unit cell 2 sample, more packing of droplets resulted in shorter pillars and shallower angled pillars than that what was seen in the unit cell 3 samples.Figure 6: Frame by frame study of four droplet deposition. Droplet frequency = 27 Hz.The lattice structures were printed at the frequencies specified in Table 3.

Droplet frequency = 27 Hz.The lattice structures were printed at the frequencies specified in Table 3. The printing of the slanted pillars was captured in high speed video using an Edgertronic SC2+ video camera at a frame rate of 5000 frames per second. These videos were studied to understand the landing and deformation of the droplets while printing (Figure 6). From the frame by frame image study, it appeared that the droplets were spreading to attain thin disk like shapes. In some cases, droplet oscillations after the deposition were also observed.ConclusionsA set of experiments was conducted in order to assess the feasibility of producing engineered lattice structures without the need for support structures using drop-on-demand magnetohydrodynamic liquid metal droplet jetting. It was found that a droplet jetting frequency ranging from 12-27 Hz was able to produce mechanically sound inclined struts used in the lattice763structures.

It was found that a droplet jetting frequency ranging from 12-27 Hz was able to produce mechanically sound inclined struts used in the lattice763structures. A droplet overlap fraction ranging from 0.5 to 1 produced struts of varying incline angles with respect to the substrate.The ultimate aim of this research is to produce solid-skinned mechanical components with internal engineered lattice structures. Ongoing research with this technique is being conducted to print struts having arbitrarily large strut diameters. Mechanical properties of the lattice structures are also being studied. In the future, solid-skinned parts with embedded lattice structures will be printed.References[1] Gockel, J., Beuth, J. and Taminger, K., 2014. Integrated control of solidification microstructure and melt pool dimensions in electron beam wire feed additive manufacturing of Ti-6Al- 4V. Additive Manufacturing, 1, pp.119-126.[2] Gu, J., Cong, B., Ding, J., Williams, S.W.

brication of Titanium Aluminide Matrix Composites by Laser Engineered Net ShapingWeiping Liu and John DuPontDepartment of Materials Science and Engineering, Lehigh University, Bethlehem, PA 18015AbstractTiAl-based titanium aluminide alloys and their composites reinforced with ceramic particles are considered to be important candidate materials for high temperature structural applications. Laser Engineered Net Shaping (LENS) is a layered manufacturing process, which involves laser processing fine powders into three-dimensional components directly from a CAD design. In this work, the LENS process has been employed to fabricate carbide particle reinforced titanium aluminide matrix composites using TiC and gas-atomized Ti-48Al-2Cr-2Nb powders as the feedstock materials. The composites deposited by the LENS process exhibited a susceptibility to solid-state cracking due to the generated high thermal stresses.

The composites deposited by the LENS process exhibited a susceptibility to solid-state cracking due to the generated high thermal stresses. The microstructures of the laser-deposited monolithic and composite titanium aluminide materials were characterized using light optical microscopy and XRD techniques. Effects of the LENS processing parameters on the cracking susceptibility and microstructure were studied. Crack-free deposits can be fabricated by preheating the substrate to 450~500°C during LENS processing. The fabricated composite deposits exhibit a hardness of more than twice the value of the Ti-6Al-4V alloy.Keywords: LENS processing; Titanium aluminides; Composites; Direct laser fabrication1. IntroductionTitanium aluminide alloys based on TiAl and their composites reinforced with ceramic particles have been considered and are being developed for high temperature applications in the automotive, aerospace and power generation industries [1].

The TiAl-based alloys have several advantages over the conventional titanium alloys, such as higher elasticity modulus, lower density, better mechanical properties at elevated temperatures and higher oxidation resistance due to the formation of a surface passivated alumina layer [2]. The intermetallic-matrix composites (IMCs) reinforced with ceramic particles generally possess even higher specific strength, specific stiffness, increased creep strength, improved toughness and high temperature strength retention. Discontinuously reinforced intermetallic-matrix composites have been produced by a variety of techniques, most of which rely on a powder processing route including mechanical alloying and combustion reaction synthesis or reactive sintering technique, cryomilling, reactive hot isostatic pressing and XD synthesis [3].

Despite a number of advantages, fabrication of IMCs by powder metallurgical processing generally requires the use of dies for pressure-aided densification, which limits its applications. Laser Engineered Net Shaping (LENS) is a solid freeform fabrication (SFF) process, which involves laser processing fine powders into fully-dense three-dimensional components directly from a CAD design. A variety of metals and alloys have been deposited by LENS processing, such as H13 steel, 316 stainless steel, nickel-base superalloys and titanium alloys [4-7]. However, little work has been done on processing titanium aluminide intermetallics by LENS although a few papers were published on microstructural characterization of laser-deposited TiAl alloys124[8-9]. So far, no effort has been made to fabricate aluminide intermetallic matrix composites by LENS.

So far, no effort has been made to fabricate aluminide intermetallic matrix composites by LENS. In this work, a feasibility study has been conducted to fabricate carbide particle reinforced titanium aluminide matrix composites by LENS processing using TiC and gas-atomized Ti-48Al-2Cr-2Nb powders as the feedstock materials. Two types of applications are expected for the LENS processing of ceramic particle reinforced titanium aluminide matrix composites. One is the fabrication of deposits on a titanium substrate alloy in order to improve its resistance to high temperature and/or other surface properties such as wear resistance. The other is the rapid production of a three-dimensional component of the intermetallic matrix composite.2. Experimental ProcedureThe materials used in this investigation were gas-atomized pre-alloyed Ti-48Al-2Cr-2Nb (in at.%) powder obtained from Crucible Research, and TiC powder (purity 99.9%) from Atlantic Equipment Engineers.

All the powders have a mesh size of –100/+325 (particle sizes between 45 and 150 µm). A hot-rolled Ti-6Al-4V plate of 3.175 mm in thickness was used as the substrate material, which had a composition of 6.1 Al, 4.3 V (in wt.%) and was obtained from Trans-World Alloys. It is a typical (α+β) Ti alloy of wide industrial applications. The microstructure of the Ti-6Al-4V substrate is composed of equiaxed and elongated α and intergranular β phases. The substrate was ground with 320 grit SiC paper, and then degreased using acetone and ethanol before deposition. In the present work, an Optomec LENS 750 machine was used to deposit the samples, which has a size of 12.7 mm × 12.7 mm × 8 layers. The LENS machine consists of a 750 W Nd:YAG laser, a 4-nozzle coaxial powder feed system, a controlled environment glove-box and a motion control system [10].

The LENS machine consists of a 750 W Nd:YAG laser, a 4-nozzle coaxial powder feed system, a controlled environment glove-box and a motion control system [10]. For fabrication of the composite deposits, 10 and 20 vol% TiC powder respectively, were added to the Ti-48Al-2Cr-2Nb powder and mixed thoroughly before being loaded into the powder feeder. Since TiC has a similar density (4.9 g/cm3) to that of (4.5 g/cm3), homogeneous mixing can be achieved. Monolithic Ti-48Al-2Cr-2Nb Ti-48Al-2Cr-2Nb deposits were also produced for comparison. Different laser powers (from 170 to 340 W) and traverse speeds (from 4.2 to 16.9 mm/s) were used in the experiments. The powder feed rate varied from 2.4 to 3.5 g/min according to the traverse speed. The layer thickness, hatch spacing, and stand-off distance were set at 0.254 mm, 0.381 mm and 152.4 mm respectively. The oxygen level in the glove box was kept below 6 ppm during processing.

The layer thickness, hatch spacing, and stand-off distance were set at 0.254 mm, 0.381 mm and 152.4 mm respectively. The oxygen level in the glove box was kept below 6 ppm during processing. Light optical microscopy, electronic probe micro-analysis (EPMA) and XRD were used for microstructure and phase analysis. The surfaces of cracks were examined with a JEOL 6300 SEM. Microhardness measurements were conducted on selected specimens using a Vickers indenter under a 300 gram load.3. Results and Discussion3.1 Microstructure and phase analysisFig. 1 shows the microstructure of a Ti-48Al-2Cr-2Nb deposit processed with a laser power 300 W and 6.4 mm/s traverse speed. The microstructure of as-deposited monolithic Ti-48Al-2Cr-2Nb consists of massively transformed γ phase (dark contrast) and a small amount of α2 phase (light contrast). The α2 phase was mostly observed in the first layer adjacent to the125substrate (as shown in Fig.

The α2 phase was mostly observed in the first layer adjacent to the125substrate (as shown in Fig. 1a) and at the bottom of a layer in a vertical cross-section or at the inter-track region in a horizontal cross-section. This is because the cooling rate is much higher at these regions under the heat conduction conditions during processing, such that the massive transformation to γ phase is incomplete and the high-temperature α phase is ordered to α2 phase. As is recently known [11-12], the α→γ massive transformation occurs when the cooling rate is high enough for the growth of lamellae or Widmanstatten laths to be suppressed, but not so high(a)(b)Fig.

The latter transformation should be avoided or kept at a minimum through a proper choice of processing parameters since the Al-supersaturated α2 (Ti3Al) phase is extremely brittle. As shown in Fig. 2(a), the result of X-ray diffraction conducted on the surface perpendicular to the build direction confirms that the sample contains γ and α2 phases, and the intensity of diffractive peaks for α2 is very weak due to a small volume fraction.126A C B80(a) (a)(b) (b)Fig. 3 Microstructure of Ti-48Al-2Cr-2Nb+20v%TiC composite deposit (as-polished) processed with a 300 W laser power and 8.5 mm/s traverse speed: (a) light micrograph; (b) SEM micrograph.The micrograph in Fig. 3 shows that the microstructure of Ti-48Al-2Cr-2Nb+20v%TiC deposit fabricated using a 300 W laser power and 8.5 mm/s traverse speed consists of unmelted TiC particles, dendritic and needle-like phases (Fig. 3b at higher magnification) distributed in the titanium-aluminide matrix. The XRD result (Fig.

3b at higher magnification) distributed in the titanium-aluminide matrix. The XRD result (Fig. 2b) indicates that TiC and Ti2AlC carbides are present in the sample in addition to γ and α2 matrix phases. Compared with Fig. 2a, the relativeintensity of diffractive peaks for α2 in the composite deposit is indicating enhanced, significantly an increase fraction. The the volume in electronic probe micro-analysis conducted on the sample revealed that the dendritic phase is mainly composed of Ti and C (about 17.7 wt% carbon) with a small amount of Al and Nb. Thus, the dendritic phase is the primary TiC formed from the melt upon cooling. the Under substrate and powders absorb energy from the beam and form a melt pool. It is known that the laser absorption coefficient (α) of a material is a function of temperature, itslaser beamirradiation,Fig.

It is known that the laser absorption coefficient (α) of a material is a function of temperature, itslaser beamirradiation,Fig. 4 A partial isothermal section of the Ti-Al-C ternary system at 1050°C [14].electrical resistivity and the laser’s wavelength [13], as given by: α(T)=0.365{ρ20[1+ρ(T-293)]/λ}0.5−0.0667{ρ20[1+ρ(T-293)]/λ}+0.006{ρ20[1+ρ(T-293)]/λ}1.5 where, ρ and ρ20 are the material’s electric resistivity (in Ωcm) at temperature T and 20°C respectively, λ is the incident laser’s wavelength, and T is temperature (in K). The TiC powder127particles have a much larger value of electric resistivity than the titanium aluminide particles, and therefore a higher absorption coefficient. In addition, the heat conductivity of TiC is lower. Consequently, in spite of its high melting point (3067°C) partial melting of the TiC powder particles occurred in the melt pool.

In addition, the heat conductivity of TiC is lower. Consequently, in spite of its high melting point (3067°C) partial melting of the TiC powder particles occurred in the melt pool. As indicated in Fig.3, the particle size of unmelted TiC is much smaller than the starting particle size (45 to150 µm). It was also found that the amount of melted TiC is dependent on the processing parameters used. The amount of melted TiC increased with the increasing laser incident energy or heat input (incident laser power over traverse speed). The melted TiC mixing with the molten Ti-48Al-2Cr-2Nb formed a Ti-Al-C-Cr-Nb liquid in the melt pool. The amount of melted TiC was estimated to be approximately 50% of the total amount through image analysis for the sample shown in Fig. 3(a). The composition of the liquid formed in the melt pool was then calculated to be Ti-40.9Al-6.7C-1.8Cr-1.8Nb.

3(a). The composition of the liquid formed in the melt pool was then calculated to be Ti-40.9Al-6.7C-1.8Cr-1.8Nb. According to a partial isothermal section of the Ti-Al-C ternary system at 1050°C [14], the microstructure in equilibrium for this composition consists of H (Ti2AlC), γ and α2 phases without the presence of TiC phase, as shown in Fig. 4. The discrepancy from the observed microstructure can be(a)(b)Fig. 5 Cracks observed Ti-48Al-2Cr-2Nb+20v%TiC (250 W, 6.4 mm/s) deposits.in (a) Ti-48Al-2Cr-2Nb (300 W, 8.5 mm/s) and (b)Fig. 6 SEM fractographs at different magnifications of crack surfaces observed in a Ti-48Al-2Cr-2Nb deposit (300W, 10.6 mm/s).128explained from the solidification path under rapid cooling conditions.

6 SEM fractographs at different magnifications of crack surfaces observed in a Ti-48Al-2Cr-2Nb deposit (300W, 10.6 mm/s).128explained from the solidification path under rapid cooling conditions. According to a partial TiAl-TiC section of Ti-Al-C system proposed by Zakharov et al [15], at a carbon content higher than about 2 at.% primary TiC1-x solidifies first from the melt at high temperatures (>2200°C), then a peritectic reaction TiC1-x + L→ Ti2AlC takes place at about 2200°C. Below this temperature, the Ti2AlC ternary carbide can also precipitate from the liquid due to a decrease in carbon content of the liquid composition. Another peritectic reaction L + Ti2AlC → TiAl can occur later. According to Cam et al [14], a reaction L + β → α + Ti2AlC may take place at the final stage. On the basis of the available information, the phase transformation process may be described as L → TiC+L → TiC+Ti2AlC+L → TiC+Ti2AlC+β+L → TiC+Ti2AlC+α → TiC+Ti2AlC+γ(TiAl)+α2(Ti3Al).

On the basis of the available information, the phase transformation process may be described as L → TiC+L → TiC+Ti2AlC+L → TiC+Ti2AlC+β+L → TiC+Ti2AlC+α → TiC+Ti2AlC+γ(TiAl)+α2(Ti3Al). The observed increase of α2 volume fraction in the composites from the XRD results can be attributed to the fact that carbon atoms in solid solution promote the stability of the high-temperature α phase [16]. It can also be found from the Ti-Al-C phase diagram in Fig. 4 that the solubility of carbon in α2 phase is much higher than that in γ phase.3.2 Cracking problem and its prevention0.7The initially encountered in depositing both the monolithic and composite titanium aluminides by LENS is the occurrence of cracking in the deposits. In most cases, the observed cracks originated from the edge or surface of the samples with the crack plane being perpendicular to the substrate (Fig. 5).

In most cases, the observed cracks originated from the edge or surface of the samples with the crack plane being perpendicular to the substrate (Fig. 5). But cracking was sometimes also seen to occur at the interface between the deposit and the substrate, as shown in Fig. 5a. representative SEM Fig. 6 shows fractographs of in Ti-48Al-2Cr-2Nb deposits. As can be seen, fracture occurred principally by transgranular cleavage, which is typical of failure in titaniumprincipalproblemthe the cracksurfaces0.6)m m/ r e b m u n (0.50.4y c n e u q e r f0.30.2k c a r C0.10.010203040506070Incident energy, P/V (J/mm)Fig. 7 Dependence of cracking frequency on the incident laser energy for the composite deposits.(a)(b)Fig. 8 Effectiveness of preventing cracking in Ti-48Al-2Cr-2Nb+TiC composite deposits by heating the substrate during processing: (a) without preheating; (b) heating to 450°C.12980aluminide intermetallics.

The cracking is therefore a type of solid-state cracking. It was also found that the number of cracks (crack frequency) in the vertical cross-section of the cracked deposits varied according to the processing parameters used (the incident laser power (P) and traverse speed (V)). As shown in Fig. 7, crack frequency decreased as the incident energy (P/V) of laser beam was increased. It is known from the welding research that the cooling rate in the laser deposition process is mainly dependent upon the heat input (P/V) for a given material.

The reduced cracking propensity at higher heat inputs is ascribed to the resulting lower cooling rates. It is believed that cracking was caused by a high level of thermally-induced stresses exceeding the material’s a temperature range where ductility of the material is minimal, as also indicated by Patterson et al [18] in their weldability study of a titanium aluminide alloy. A the rate higher thermally-induced stresses, leading to an increase in cracking susceptibility. As the ductility of the intemetallic alloys and composites is relatively lower than that of steels and conventional alloys, these materials are more susceptible to cracking during LENS processing. is Experiments difficult the to eliminate cracking deposits only by regulating the processing parameters although cracking can be mitigated in this way. A hot plate was used to preheat the substrate to 450~500°C during LENS processing for preventing the solid-state cracking.

Fig. 8 shows photographs of as-polished surfaces of the aluminide composites deposited without and with substrate heating during processing using the same parameters, respectively. It is therefore a very effective method for eliminating cracks in the LENS deposits.3.3 MicrohardnessThe microhardness (HV0.3) was measured from the deposit surface down to the substrate. Fig. 9 (a) and (b) show the hardness distributions for the composite deposits processed at different traverse speeds, while Fig. 9(c) is the result for a monolithic titanium aluminide deposit for comparison. The measured microhardness values for the composite are the overall bulk hardness without the inclusion of unmelted TiC particles (i.e. the microhardness data are not used when the indenter hit unmelted TiC particles).

the microhardness data are not used when the indenter hit unmelted TiC particles). The microhardness of the unmelted TiC particle the was measured Ti-48Al-2Cr-2Nb+20v%TiC composite deposits, the one processed with a slower speed is harder with an average hardness of 700 VHN, which is approximately 2.2 times that of the substrate (320 VHN). This is because the deposit contains more re-solidified TiC and Ti2AlC carbides resulting from a larger amount of melted TiC when a slower traverse speed is used at a same laser power. The large scatter of the measured hardness in the deposit was due to the inhomogeneous distributions of the dendritic and needle-like carbides together with γ and α2 matrix phases of different hardness. The monolithic titanium aluminide deposit has an average hardness value of 380 VHN, which is similar to that obtained by the conventional processes. It is noted that in all cases the hardness in the layer adjacent to the substrate is the highest.

It is noted that in all cases the hardness in the layer adjacent to the substrate is the highest. Based on microstructural analysis, it is found that the volume fraction of α2 phase is higher in the first layer as a result of the dilution of the substrate combined with a higher cooling rate. The heat-affected zone (HAZ) has a slightly higher hardness than that of the substrate as the HAZ has an acicular martensite structure (the supersaturated α’-phase). The high hardness of the composite deposits is expected to exhibit an excellent wear resistance. In addition, through compositional and processing optimization, the carbide reinforced γ-TiAl matrix composites are also expected to possess improved strength and fracture toughness. Further work is needed in these aspects.to be 2240~2390 Vickers hardness number(VHN). For4.

Further work is needed in these aspects.to be 2240~2390 Vickers hardness number(VHN). For4. ConclusionsA feasibility study has been conducted to fabricate carbide-particle-reinforced TiAl matrix composites using TiC and Ti-48Al-2Cr-2Nb powders as the feedstock materials by LENS processing. Both the monolithic and composite titanium aluminide deposits exhibited a susceptibility to solid-state cracking due to the high thermal stresses generated from the LENS processing. Crack-free deposits can be fabricated by preheating the substrate during LENS processing. The microstructure of fabricated composite deposits consists of unmelted TiC particles with resolidified TiC and Ti2AlC carbides distributed in γ and α2 matrix phases. The microstructure of monolithic Ti-48Al-2Cr-2Nb deposits is composed of massively transformed γ phase and a small amount of α2 phase.

At present, most of the powder manufacturing techniques are being focused on layering powder with different powder blend compositions with Z gradients (graded in direction of layer build). Although, there are a few researchers working on multi powder feeder and deposition system, the study of laser fusion of the deposited powder (by a powder deposition system) is minimum or not known to date. Consequently, the manufacturing of functionally graded structures is still geometry limited. This work was focused on the manufacturing of X-graded (graded along the powder bed plane) specimens with H13 tool steel and Cu mix. Five bimodal powder blends were used with a multi-container feed hopper to spread powder layers for the selective laser fusion of the powder. The powder was fused using a high power Nd:YAG pulsed laser using a specific scanning strategy to reduce porosity. Specimens were produced with graded Cu the H13 matrix.

The powder was fused using a high power Nd:YAG pulsed laser using a specific scanning strategy to reduce porosity. Specimens were produced with graded Cu the H13 matrix. The specimens were analysed for dimensional accuracy, within microstructure, porosity, cracks and micro hardness of the FGM.Keywords: functionally graded materials, laser fusionIntroductionFunctionally Graded Materials (FGM) has been in the focus of researchers since the 80’s (Suresh & Mortensen; 1998). The main advantage of exploring the functionally graded manufacturing techniques and materials combinations is the possibility of producing highly optimised parts for high performance applications. Unfortunately, most of the graded materials and processes are geometry dependant and engineers are not able to use these materials easily.Lately a number of researchers are working on the application of layered manufacturing (LM) technologies to produce FGM parts.

As these technologies can virtually manufacture free forms without any geometry limitations, it would be the best possible approach to make the FGM parts fabrication successful. Furthermore, as the LM technology is an additive process, this would make the deposition of the powder in the required composition easy and flexible.The use of layered technologies to produce FGM parts is limited to the development of Computer-aided Design (CAD) and Finite Element Analysis (FEA) data, Computer-aided Manufacturing (CAM) processing and Local Composition Control (LCC) of materials. The CAD and FEA are important tools as these aid the engineer in deciding the composition of material in various regions of the part and the transition between the materials. Also, CAD187data for FGM parts should be interchangeable with FEA for optimum engineering design. The CAM must be able to read the CAD data with its different materials and process it for each layer/step during the fabrication process.

The CAM must be able to read the CAD data with its different materials and process it for each layer/step during the fabrication process. Finally, LCC is a new and innovative devices that can deliver different compositions of materials to a layer according to the CAM instructions. However, at present the LCC devices are in early stages of research and development.This work uses a simple X-graded powder deposition system to deposit graded layers to produce FGM parts. However, these parts are limited to X-grading along the powder bed surface for selective laser fusion with high power Nd:YAG pulsed laser. At this stage, the system is used for parts fabrication in order to optimise laser parameters which improve the mechanical properties and minimizes porosity and cracks. The paper presents the results obtained from the initial work.

The paper presents the results obtained from the initial work. The microstructure of the material, dimensional accuracy, microhardness, porosity and cracks were analysed.Background Functionally graded materials are given importance because of the possibilities of combining distinctive materials to add new functions to the same component. Also, designers could optimise form and function in their new projects improving the performance of the products. According to Suresh & Mortensen (1998), FGM could be used to manufacture components with engineered gradual transitions in microstructure and/or composition. This gradual transition results in different functional performance requirements that vary with location within the component.Nowadays, the manufacturing of the FGM parts is based on constructive and transport-based processes (Suresh & Mortensen; 1998). Constructive processes are based mainly on powder densification/sintering techniques and coating processes.

Constructive processes are based mainly on powder densification/sintering techniques and coating processes. Transport-based processes are those associated with mass transport, thermal process, settling and centrifugal processes and infiltration and macro segregation processes. Many pre-forms and simple geometries are being produced using these processes. However, there is still lack of “smart” processes (such as free-form fabrication) that could deliver complex parts to unleash all the possible applications for FGM.In order to improve and produce complex parts using the FGM technique, many researchers from layered manufacturing technologies, often referenced as Rapid Prototyping and Manufacturing (RP&M), are working on this area. Selective laser sintering (SLS) works with fine layers of powder deposited over a platform. Each layer is selectively sintered in some regions using a laser beam.

Selective laser sintering (SLS) works with fine layers of powder deposited over a platform. Each layer is selectively sintered in some regions using a laser beam. After the layer is completed, the platform is lowered and a new powder layer is spread over the previous one. The laser beam sinters the new layer adding it to the previous one, and the process is repeated until a complete part is built. SLS uses a metallic powder coated with a thermoplastic binder to give shape to the part through the melting of the binder. Subsequent steps are binder removal, pre-sintering, sintering and copper/bronze infiltration. Currently, three kinds of materials that can be used; LaserForm ST 200 (54% 420 stainless steel and 46% bronze), LaserForm ST 100 (60% 420 stainless steel and 40% bronze) and LaserForm A6 (Steel, not much information available) (3D System 2004).

Direct metal laser sintering (DMLS) is based on the same process of SLS but it uses the power of the laser to sinter in loco the powder, without the necessity for special post- processing in an oven to increase the density of the part. Nevertheless, the parts present high porosity and lower mechanical properties compared to those obtained by selective laser sintering (Storch et al., 2003). There are two materials currently available for this process; DirectMetal 20 & 50µm (nickel, bronze and copper phosphate mix) and DirectSteel 20 &18850µm (steel alloy, details are not available). Laser cladding works by coaxially blowing powder with shield gas over a high power density laser spot and subsequently fusing the powder. Laser cladding is successful in using more materials than other two technologies (Atwood et al., 1998; Ensz, et al., 2002; Lü et al., 2001). This technique can be used for building fully dense parts.

Laser cladding is successful in using more materials than other two technologies (Atwood et al., 1998; Ensz, et al., 2002; Lü et al., 2001). This technique can be used for building fully dense parts. Unfortunately, parts made by this method present some geometrical limitations and low tolerances. Usually it is necessary to perform milling operations after building the part but many advances have been made in recent years (Lü et al, 2001).These additive free-form manufacturing technologies can produce free-form parts of any complex geometry. One of the main issues to produce FGM in RP&M machines is the Local Composition Control (LCC). A couple of works such as Kumar et al (2004) and Yang et al. (2004) are investigating the use of capillary tubes to control the deposition of powder. Ensz et al (2002) have been studying the optimization of two powder flows into the LENS process (Laser Engineering Net-Shaping) to build gradients from H13 to M300 steel alloys.

Ensz et al (2002) have been studying the optimization of two powder flows into the LENS process (Laser Engineering Net-Shaping) to build gradients from H13 to M300 steel alloys. Jepson et al (2003) have been working on the development of a selective laser machine to build FGM layers using a polymeric binder to give shape to the parts prior to de-bind and to sinter in the furnaces. Cho et al (2002) have investigated the LCC for the 3D Printing process by obtaining data for a model from FEA and Voxel space geometries.Various researchers have investigated the materials that can be processed by this method. In Liverpool University, UK fully dense stainless steel specimens have been made using nanosecond pulsed Nd:YAG lasers by Direct Metal Laser Re-melting (DMLR), a variation of SLS process (Morgan et al, 2001). Also porous thin wallets structures with copper powder melting have been investigated through the same method (Pogson et al, 2003).

Also porous thin wallets structures with copper powder melting have been investigated through the same method (Pogson et al, 2003). It was found that an insufficient melting of the copper occurred in the samples. Also, continuous wave mode achieved better results than the pulsed laser mode in the fusion of the Cu powder. The production of dense H10 tool steel parts and graded H10 with tungsten carbide (WC) was evaluated by Su (2002). FGM structures of H10 with percentages up to 20% of WC had increased the micro hardness of the H10 matrix. Leong et al (2002) have produced copper matrix composites through laser sintering with a CO2 laser. Porous parts with Ti-B4C-Cu were made. The addition of nickel (Ti-B4C-Cu-Ni) improved the wetting and diffusion between the particles and decreased the porosity.

Porous parts with Ti-B4C-Cu were made. The addition of nickel (Ti-B4C-Cu-Ni) improved the wetting and diffusion between the particles and decreased the porosity. Also nickel, aluminium and stainless steel powder fusion was investigated by Abe et al (2001) using two sources of laser energy, CO2 and Nd:YAG, with strategies of pre-heat and re-heat the material. Using dual scanning with Nd:YAG laser melting and CO2 laser re-heating improved the materials ductility.MethodologyAt present, a simple and manual X-graded powder deposition system, Figure 1, was designed to evaluate parts produced and study the processing issues. In order to obtain the X-graded regions, a powder feed hopper was built with eight compartments for the mixture of powders. The pre-mixture powders were placed in the compartments of the hopper according to the desirable composition design. The hopper slid over a platform spreading the powder mixtures in trails over a substrate placed on the platform.

The hopper slid over a platform spreading the powder mixtures in trails over a substrate placed on the platform. A high power fibre-optic delivered Nd: YAG laser was used to fuse the layer. Argon gas assist was used in order to avoid the oxidation of the material. After completion of the layer, the platform was lowered and a new powder layer was spread over the previous fused layer. The laser fuses the new layer, and the process is repeated until the part is completed.189Figure 1. A schematic diagram of x-graded de-coaterThe H13/Cu powders were mixed in five different compositions. The powder blends were tailored to achieve the particle size ratio of 7:1 to increase the apparent powder density. The powder mixture gradients from H13 to Cu were made by discrete increments of 12.5% until reaching 50% of Cu in H13 powder. The powder was sieved and particles in the range of 106 to 150µm for H13 and Cu were mixed with particles below 38µm and 22µm of H13 and Cu respectively.

The powder was sieved and particles in the range of 106 to 150µm for H13 and Cu were mixed with particles below 38µm and 22µm of H13 and Cu respectively. The H13 tool steel powder was spherical and the Cu powder was more irregular and agglomerated. The H13 tool steel alloy composition was, in percentage weights: Fe 91.2%, C 0.4%, Cr 5.2%, Mn 0.4%, Si 1%, V 1% and Ni max. 1% and the copper was OFHC Cu (Osprey, 2003). Figure 2 shows images obtained in a Scanning Electron Microscope (SEM) from three of the five used mixtures/blends (H13, H13+25%Cu, H13+50%Cu).Figure 2. SEM images from the powder blends (a) H13, (b) H13 + 25%Cu and (c) H13 + 50% Cu. Note: Only large particles are indicated.190The powder mixture of H13/Cu was filled in each compartment of the hopper (see Figure 1) in sequence by increasing Cu percentage to the 50% in the middle compartment. The powder layer thickness was approximately 300µm and the substrate used was mild steel plate.

The powder layer thickness was approximately 300µm and the substrate used was mild steel plate. From the work conducted previously at Loughborough University (Su, W., 2002), the refill laser scanning strategy (Figure 3) has been identified as an accurate building strategy. The strategy consisted of alternate scanning vectors followed by a refill of powder and scanning the vector- gaps. The use of this pattern, which is shown in Figure 3, is better explained in Su (2002). The pattern accounts for the shrinkage and contraction of powder during the laser processing. Some powder gets sucked to the melt pool due to capillary action and a fraction of material is lost due to spatter effect. This leaves the scanned layers un-even and voids areas that results in a pore during the fabrication of the whole cube. This effect can be minimised by the alternate scanning strategy and refill in between the scanning process.Figure 3.

This effect can be minimised by the alternate scanning strategy and refill in between the scanning process.Figure 3. The refill scanning strategyThe specimens were built using a high power Nd:YAG pulsed laser (GSI Lumonics JK701H). The process parameters were optimised in the previous work (Su, W., 2002). The pulse energy used was 10J, repetition rate of 10Hz, pulse width of 5ms and the feed rate of 3.33mm/s. The laser spot diameter was 0.8mm. The average laser power was 100W with an energy density of 37.5J/mm2 and the peak power of 2kW. The specimens were fused under room temperature and under a localized gas assist using argon to reduce the oxidation.Specimens were built at 40x6x4mm dimension. The specimens were cross-sectioned, mounted, polished and etched with 2% of nitric acid solution. The microstructure and the microhardness analyses were performed. The dimensional accuracy analysis was performed on 10x10 mm square specimens.

The microstructure and the microhardness analyses were performed. The dimensional accuracy analysis was performed on 10x10 mm square specimens. The microhardness test was performed along the X gradient with the spacing of 1mm using Vickers hardness system.ResultsThe purpose-built X-graded recoater system’s performance was found to be good. The composition was distributed discretely over the powder bed. Figure 4, shows the distribution of the powder mixture and the fused layer of a specimen.191Figure 4. Image of the laser fused X-graded layer and the material composition.The surface quality was similar despite the variation in the H13/Cu compositions. The specimen’s top layer was rough and irregular. The surface could be re-melted by using lower laser intensity. However, this is not performed in this work as the focus is to optimise the parameter and achieve the optimal mechanical and microstructure properties.

However, this is not performed in this work as the focus is to optimise the parameter and achieve the optimal mechanical and microstructure properties. It also can be seen that there is no distinct boundaries between the various compositions of H13/Cu mixtures.The dimensional accuracy measurements were performed on six specimens. The average error obtained in the X and Y directions are 0.8% and 1.3% respectively and the standard deviation was 0.5. This results show that the process parameter used produces accurate parts in those directions. However, as the powder layer deposition is partially automated and powder layer thickness is not accurately controllable, the average error in the z direction was found to be - 6.7% with the standard deviation of 1.2.Figure 5 shows one of the specimens and the longitudinal cross-section image. The cross sections along the specimens indicated the presence of pores in all formed alloys.

The cross sections along the specimens indicated the presence of pores in all formed alloys. Cracks were strongly present in the 12.5, 25 and 37.5%Cu blends. These cracks could have been formed due to the high cooling rate achieved by the Cu addition to the H13. The formation of the cracks was found to be similar to the cracks in laser welded parts. These cracks could be formed in the middle of the laser beam and in the heat affected zone (HAZ).Figure 6 shows the cracks and porosity for the H13 and 12.5%Cu region. The cross section of the specimens showed relatively low porosity. The porosity level was measured and was found to be 2.35% at pure H13 region. In the region of H13+25%Cu and H13+50% Cu the level of porosity increased tremendously to 2.89 and 6.80% respectively. However, the percentages of porosity found in these specimens are considered to be low compared to the conventional powder metallurgy (German, 1994).192Figure 5.

However, the percentages of porosity found in these specimens are considered to be low compared to the conventional powder metallurgy (German, 1994).192Figure 5. The overall view of the specimen (top) and the longitudinal cross-section of fused part with various H13/Cu mixtures (bottom).Figure 6. Cracks and pores in the 12.5%Cu section.The pure H13 formed fine dendrites due to the effect of laser fusion of the material above the supercritical liquidus line and the rapid cooling. The alloying elements were concentrated in the interdendrite spaces (see Figure 7a). The dendrites can be formed by austenite or martensite depending on the region of the part. According to Brooks et. al. (1999) the re- melting and re-heating cyclic effects of the laser iteration with the addition of new layers cause this microstructure change.

According to Brooks et. al. (1999) the re- melting and re-heating cyclic effects of the laser iteration with the addition of new layers cause this microstructure change. With the addition of the Cu, a new alloy formed and the microstructure remains the same, however with the segregation of the Cu in the interdendrite spaces (see Figure 7b).193(a) Pure H13 (mag. 1.75k X)(b) H13+25%Cu (mag. 1.50k X)Figure 7. SEM picture of pure H13 and H13+25%Cu microstructures.With the higher percentages of Cu, in the region of 37.5% and 50% composition, a two phase material was characterized, especially in the 50% Cu region. Figure 8 shows the SEM picture where the two phase material was clearly perceptible in the 50% Cu region. Part of the Cu segregated forming isolated areas of Cu. The other part of the Cu was found in the interdendrite spaces as in the other regions. Spherical structures of the H13 and Cu could be found dispersed inside pure Cu region.

The other part of the Cu was found in the interdendrite spaces as in the other regions. Spherical structures of the H13 and Cu could be found dispersed inside pure Cu region. X-ray diffraction analysis showed that these particles were not from non-fused H13 particles and contained approximately 23% of Cu.Figure 8. SEM picture of H13+50%Cu region (mag. 500 X).The microhardness taken along the X-graded specimens shows the variation in hardness along the various regions. Figure 9 shows the microhardness results obtained from a single specimen. The graph shows that each region of the powder composition is 4mm long. The194variation of the hardness shows that the hardness reduces by increasing the Cu percentage. The variation of the hardness was observed to be high at H13+50%Cu mix region. This is mainly due to the two phase region formation as observed in the microstructure analysis.

The variation of the hardness was observed to be high at H13+50%Cu mix region. This is mainly due to the two phase region formation as observed in the microstructure analysis. From this observation, the saturation of Cu in the H13 matrix is between 37.5 and 50%.Microhardness along the X graded specimen800H13+12.5%Cu+25%Cu+37.5%Cu+50%Cu700s r e k c i v600500s s e n e d r a H4003002001000048121620Position [mm]H13 Tool steel Copper (hardned)Figure 9. Microhardness along the X-graded specimen.ConclusionsProducing FGM parts using direct metal laser melting is a complex task as it involves many areas of expertise. The work being carried out at Loughborough University is promising. The knowledge in laser processing and the associated process issues could be well utilised in producing complex FGM parts.

The work being carried out at Loughborough University is promising. The knowledge in laser processing and the associated process issues could be well utilised in producing complex FGM parts. However, there are many other areas that still need to be tackled in order to produce complex FGM shapes: powders deposition integration, dimensional accuracy and CAD/CAM interfaces.In this work, the performance of the X-graded powder deposition system was found to be satisfactory. However, the gradient is discrete and voids are limited by the powder flow ability. Smaller particles could be used but also can reduce the flow of the powder. These issues could be solved with vibratory hopper and other mechanisms as being investigated by other researchers.The porosity found in the FGM specimens was considered low. However, the cross-section image shows the formation of cracks in the H13 and Cu mixed parts.

However, the cross-section image shows the formation of cracks in the H13 and Cu mixed parts. It is necessary to optimize the laser processing parameters to moderate the cooling rate of the material and hence reduce or eliminate the crack formation. Future work will involve in determining the best laser processing parameters for each blend according to the Cu percentages. The use of195pulse shapinging and appropriate shield atmosphere and preheating could also help in reducing the cracks. The two phase material formation could be expected as the percentages of added Cu are high. The saturation of the Cu in isolated areas is evident from the variation in the microhardness results. This saturation is high in the region where the Cu percentage is between 37.5% and 50%.The surface quality and dimensional accuracy is not the major focus of this work. However, the dimensional accuracy in the x-y directions was found to be accurate, 0.8 and 1.3 percent of error respectively.

However, the dimensional accuracy in the x-y directions was found to be accurate, 0.8 and 1.3 percent of error respectively. At this stage, the surface quality is found to be poor, however, the parts could be machined or a laser re-melting technique could be used to smooth the surface.AcknowledgementThe authors would like to thank CAPES and CNPq, Brazil for supporting the research student and Osprey Powders, UK for the in kind contribution of powder materials.References3D Systems Website, LaserForm datasheets, http://www.3dsystems.com, Accessed on February 2004.ATWOOD, C., GRIFFITH, L.H., SCHLIENGER, E., ENSZ, M., SMUGERESKY, J., ROMERO, T., GREENE, D., RECKAWAY, D., 1998, Laser engineering net shaping (LENS): A tool for direct fabrication of metal parts. Proceedings of ICALEO '98, Orlando, FL, USA.BROOKS, J., ROBINO, C., HEADLEY, S., GRIFFITH, M., 1999, Microstructure and property optimization of LENS deposited H13 tool steel.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA FACET CLUSTER-BASED METHOD FOR ALTERNATIVE BUILD ORIENTATION GENERATION IN ADDITIVE MANUFACTURINGYicha ZHANG*, Ramy HARIK†, Wout DE BACKER†, Alain BERNARD*IRCCyN, Ecole Centrale de Nantes, Nantes, France †University of South Carolina, Columbia, South Carolina, USAAbstractBuild orientation determination is an important pre-processing step in Additive Manufacturing. To identify an optimal build orientation, there are two main tasks, generating a set of alternative orientations and evaluating these alternatives with pre-set criteria. To solve the first task, currently there are two categories of methods, exhaustive computing and continuous surface decomposition.

To solve the first task, currently there are two categories of methods, exhaustive computing and continuous surface decomposition. However, for exhaustive computing methods, the infinite original alternative orientation space is an obstacle, especially when considering multiple objectives. While the other type of methods have difficulty on surface separation and shape boundary recognition when facing complex CAD models. To tackle of these obstacles, this paper introduces a new method applying a statistical tool to form facet clusters for decomposing an STL model in a discrete way. The formed facet clusters can be used to efficiently generate meaningful alternative build orientations and can also be used to predict surface quality distribution over a part model for further process planning or design iteration.IntroductionAdditive Manufacturing (AM) has become a popular processing technology recently in many application domains [1].

Using a layer-by-layer additive building way, it can deal with parts with extremely complex geometries and internal structures. This brings it a great advantage for processing high-value and high- customized parts [2]. Meanwhile, the design freedom has been increased due to the decreased processing constrains on geometry [2-4]. As a digital manufacturing, AM is totally data-driven, which makes it seem that inputting a digitized CAD model can result to a final prototype or component in a fully automated way by only pushing a start bottom in the AM machine. However, except geometric constraints, in order to obtain a good production quality, there are still many other constraints that should be considered. For example, the intrinsic ‘step-effect’ and the anisotropic material properties, caused by layered building, need to be analyzed when designing or manufacturing.

For example, the intrinsic ‘step-effect’ and the anisotropic material properties, caused by layered building, need to be analyzed when designing or manufacturing. In the processing AM chain, there is a critical step to transfer a CAD data model into a processing data model which can be directly used by AM machines. This critical step is pre-processing, also called process planning [5, 6]. Among diverse process planning tasks, build orientation optimization is the most important one since it determines along which direction the materials will be added. This will affect the later slicing, scanning or tool-path generation results as well as the final production results, including build time, cost and quality.

This will affect the later slicing, scanning or tool-path generation results as well as the final production results, including build time, cost and quality. Mathematically, the build orientation determination problem can be defined as: for a given object, CAD model, in its original orientation (design orientation), to find a transformation matrix, T, such that after applying the transformation to the model, the build direction lines up along vertical axis (Z, material adding orientation). To determine the matrix, there are two main tasks: generating alternative build orientations from the infinite original solutions space and evaluating the generated alternatives with preset preference and criteria. The final optimal build orientation selected from the alternative build orientation set can be used to calculate the transformation matrix for the original model.

The final optimal build orientation selected from the alternative build orientation set can be used to calculate the transformation matrix for the original model. In real AM production practice, usually the AM machine setup and user or designer preferences are not changed or limited to some standard options but only the AM operators are shifted. Users only need to determine an optimal build orientation for a part model by using their own knowledge. However, a same given part model to be manufactured by the same AM machine, different operators may determine different optimal build orientations respectively. This would cause production stability problem, which means the production results would be different. This production stability problem would further prevent the industrialization of AM.

This would cause production stability problem, which means the production results would be different. This production stability problem would further prevent the industrialization of AM. The main reason why different optimal build 23orientations are generated by different operators would be that the alternative build orientation sets generated respectively by different operators are different. If a method can generate a relatively stable alternative orientation set, which contain the optimal orientation implicitly, different operators can identify the same optimal build orientation by using the same preset setup, preference and evaluation criteria as input of decision making. As a result, the production result would be stable. Therefore, the purpose of this paper is to introduce such kind of method for generating a stable alternative orientation set to solve the first task of build orientation optimization and also facilitate the other task.

The method is to decompose a CAD model represented in STL format into a set of facet clusters in a discrete way, which is different to currently used feature-based method using continuous decomposition manner. The obtained facet clusters will be used to generate alternative build orientation for the original non-decomposed original CAD model. To form reasonable facet clusters, a statistical non-supervised learning scheme is adopted. With the proposed method, both simple and complex CAD models can be decomposed quickly into facet clusters, which will be demonstrated by examples later in this paper. Before introducing the proposed method, the next section of this paper reviews current solutions in literature and shows the necessity of applying the proposed method.Related WorksIn literature, there are plenty of proposed solutions for build orientation optimization problem since the beginning of AM.

Some authors had already conducted comprehensive reviews on orientation optimization study [7, 8]. Hence, this section will not present another comprehensive review work of those methods but mainly discuss about the existing solutions for the first task, generating alternative build orientation set. Theoretically, an STL CAD model can be rotated freely within a build envelope of an AM machine if it can be fitted. Hence, the unlimited continuous rotation degrees around two axes except the build axis, Z, result to an infinite alternative build orientation set. However, the infinite orientation set cannot be directly used for the later evaluation task due to the computation cost. To solve this problem, researchers tried different methods to extract a finite alternative build orientation set or a reduced infinite orientation set from the theoretically infinite one.

To solve this problem, researchers tried different methods to extract a finite alternative build orientation set or a reduced infinite orientation set from the theoretically infinite one. The simplest way is to specify a small list of alternative build orientations by rotating the original part model with an interval. In the literature, some researchers applied this method to generate alternative build orientations for layout and orientation optimization [9-11]. However, this cannot guarantee the optimality since the big interval would cause the missing of potential optimal candidates. Hence, to avoid missing the potential optimal orientation, some researchers tried to specify the orientations in a small interval or even in a continuous way by using a continuous control parameter. Hur et al. [12] developed an exhaustive computation method for orientation optimization with the minimum build height as an optimization objective.

Hur et al. [12] developed an exhaustive computation method for orientation optimization with the minimum build height as an optimization objective. Alternative build orientations were generated by rotating the part model about two axis with a given rotation interval of 5 degrees. In [13], a statistical method, response surface methodology, was proposed to deal with the build style of SLA. In this method, build orientation was regarded as a continuous parameter controlled by a user-defined ‘Z-height’ range to be optimized. Therefore, the orientation generation manner in their method is also a user specification. However, due to the infinite and continuous orientation space theoretically, reducing interval or specifying a limited continuous orientation range can neither guarantee the global optimality. To deal with this problem, intensive work had been conducted to apply advanced computation method for single- or multiple- objective optimization for identifying the global orientation.

To deal with this problem, intensive work had been conducted to apply advanced computation method for single- or multiple- objective optimization for identifying the global orientation. The application of genetic algorithm is a representative [14-22]. In these works, alternative orientations were generated by using diverse genetic operators in a stochastic searching way and a fitness function, derived from orientation evaluation criteria, was designed for selection. For single-objective optimization, this kind of method can identify the optimal build orientation with an acceptable computation cost. But, for multi-objective optimization, the exhaustive computing is not efficient [8, 21]. As stated by Padhye and Deb [21], a possible route to arrive at the optimal orientation without carrying out exhaustive computation is to analyze the features of a solid object.

As stated by Padhye and Deb [21], a possible route to arrive at the optimal orientation without carrying out exhaustive computation is to analyze the features of a solid object. Therefore, using surface shape or features of the part model to generate alternative build orientations is another group of methods. In the beginning, part model’s convex hull was used to generate alternative build orientations. The planar surface or facet on the part model which are in the common plane with the part model’s convex hull is used as a base, of which the perpendicular orientation is identified as an alternative build orientation. A couple of representative works can be found in [23-26]. These methods can extract meaningful planar surfaces on a part model to generate a set of practical alternative build orientations. However, other concave surfaces may 24also imply some potential practical or even optimal build orientations.

However, other concave surfaces may 24also imply some potential practical or even optimal build orientations. To avoid missing these possibilities, some researchers focused on using the original part model to generate alternative build orientations. In [27-29], the planar surfaces on the part model were directly used to generate vertical directions as alternative build orientations. An extreme example was reported by [30], where each triangular facet of an STL model was used to generate an alternative build orientation along the facet’s normal direction. These methods can generate more alternative build orientations, especially the method of using facet normal. However, for part models with non- planar surfaces or meaningful shape features, these methods would encounter problem to identify additional orientation candidates implied by these surfaces or features.

As stated by Zhang and Bernard [8], in real practice, not all the surface shapes of a part model were treated equally but more attention would be paid to some local critical shape features, which implies the importance to analyze the shape features of part model in process planning. Hence, since the beginning, many researchers have investigated the using of shape features to generate alternative build orientations for optimization in AM. Frank and Fadel [31] defined 9 types of geometric features and associate a feature axis, which was used to indicate an alternative build orientation, for each defined type. In [32, 33], key geometric features or typical surface features of an STL model were used with associated rules to generate alternative build orientations for generating optimal orientations. Feature- based methods can generate meaningful alternative build orientations so as to capture the design intention [8, 34].

Feature- based methods can generate meaningful alternative build orientations so as to capture the design intention [8, 34]. However, in feature-based methods, the definition of surface feature or key surface features is not very clear to be practical. In addition, in these researches, each surface feature can only generate a fixed alternative orientation, which cannot meet the requirements of different AM processes. Since the constraints on a same shape feature may be different when using different AM processes for production. This requires a shape feature can generate different alternative orientations in different processing contexts. To tackle of these drawbacks, Zhang et al. [8, 35-38] proposed to defined specific shape features for AM as AM features to solve the process planning and design problems. User-defined AM features can generate a set of finite alternative build orientations by using the shape boundary and facet normal information as well as user preference.

User-defined AM features can generate a set of finite alternative build orientations by using the shape boundary and facet normal information as well as user preference. However, the main difficulty of feature-based methods is feature recognition due to the insufficient topological information of STL part models, especially for part models with complex geometries.In general, the discussed methods above can be summarized as two main groups: one is exhaustive computing and the other is continuous surface decomposition. Exhaustive computing methods means the generation of alternative orientations is similar to Enumeration method. For this group of methods, if the number of generated alternatives is small, the optimality of result cannot be guaranteed since a large section of the original solution space would not be explored.

For this group of methods, if the number of generated alternatives is small, the optimality of result cannot be guaranteed since a large section of the original solution space would not be explored. But, if the number of generated alternatives is large, the later computation of orientation evaluation would be ineffective, especially when facing multi-objective optimization. In addition, for this kind of method, much computation is spent on evaluating meaningless alternatives. Continuous surface decomposition means to decompose an STL model into a set of surface fragments (facets, planar surfaces or special shape features), which are composed of a set of adjacent flat facet triangles. This group of method uses decomposed surface fragments and associated rules to generate alternative build orientations. In this way, meaningful alternative build orientation can be generated and further computation of orientation evaluation can be saved.

In this way, meaningful alternative build orientation can be generated and further computation of orientation evaluation can be saved. However, as discussed above, how to decompose an STL model in a continuous way to form meaningful predefined shape features is difficult. Current feature recognition methods can only recognize simple shape features due to the lack of topological information in STL files [39]. Hence, the decomposition method should be improved. The purpose of this paper is to introduce a new decomposition method to fraction an STL model in a discontinuous way so as to help generate alternative build orientations. The next section of this paper presents the new method.Facet Clustering MethodCurrently, STL format is still the most widely-used in AM as a de-facto standard though some new formats, e.g. 3MF, have been developed. Hence, the proposed new decomposition method is based on this format.

3MF, have been developed. Hence, the proposed new decomposition method is based on this format. By using STL format, a 3D CAD model can be represented by a set of numbered triangular faces with normal and color information. A facet in ASCII STL format can be described by Figure 1 below.25Figure 1, representation of a facet in ASCII STLAccording to the planar characteristic of triangular facet in an STL model, it could be imaged that the best way to fabricate a single facet with AM processes is to align the facet’s normal direction along build orientation (‘Z’ direction) or make it perpendicular to the build orientation to avoid the stair-step effect caused by layered manufacturing. If the facet is placed in an inclined direction to the build orientation, the stair-step effect would appear.

If the facet is placed in an inclined direction to the build orientation, the stair-step effect would appear. Stepping further, when fabricating a full part model in STL format, to reduce the total impact of stair- step effect on the surface roughness or shape accuracy, it is better to orientate the part model in a direction under which the number of inclined facets is the smallest. Traditionally, a full STL model would be decomposed into a set of adjacent triangles to represent predefined shape features or pre-identified key surfaces. Then, a couple of orientation rules would be associated to these decomposed surface fragments to generate potential optimal build orientations to benefit these fragments. Finally, these generated orientations are used as alternative build orientations for the original full STL model.

Finally, these generated orientations are used as alternative build orientations for the original full STL model. As discussed above, one difficulty for the continuous decomposition method is the low efficiency of feature recognition due to the insufficient geometric information of STL file. Another ignorance of these methods is that they only considered to separate the STL model in a continuous way but never tried to analyze different disconnected surfaces which may have similar characteristics on the same model in a discrete way. In an STL model, different disconnected surfaces may have parallel or coaxial characteristic, which makes these surfaces have a similar production results (Figure 2).Figure 2, disconnected parallel and coaxial surface shapesIn Figure 2, if applying traditional continuous decomposition, the two pairs of surfaces would result to four independent surface features, two cylindrical and two planar surface features.

However, they would generate similar alternative build orientations for the part model. This will produce duplicate alternative orientations to waste more computation in the orientation evaluation stage. But, if the disconnected parallel planes or coaxial cylindrical surfaces are grouped as one surface feature, the duplicate alternative orientation for the part model would be reduced. Actually, in real practice, the build orientation would have similar impact on the grouped surface features. Therefore, in this paper, facet cluster is proposed to describe this kind of disconnected surface feature group. And then applying orientation rules to the facet cluster for generating potential optimal build orientations to benefit the cluster as did for traditional continuous surface features in [8, 31]. However, how to cluster facets in a reasonable way? This is the core question for this paper.

However, how to cluster facets in a reasonable way? This is the core question for this paper. To answer this question, there are26two tasks to do: one is the formation of a facet cluster by using limited geometric information of an STL model and the other is the identification of a suitable number of clusters for the STL model.To form a facet cluster, the normal information of the facet is used. In this paper, according to the characteristics of layered manufacturing, an assumption is made that the build orientation would have similar impact on those facets, which have similar normal directions, of an STL model. Based on this point, facets which have similar normal directions should be clustered as one group for production quality analysis. Here, similarity between a pair of facets is measured by the angle between their two normal directions. To compute the similarity among triangular facets of an STL model, a classical statistical clustering method, K-means, is adopted.

To compute the similarity among triangular facets of an STL model, a classical statistical clustering method, K-means, is adopted. K-means is an unsupervised learning algorithm that solves the well-known clustering problem [40, 41]. The clustering procedure follows a simple and easy way to classify a given data set through a certain number of pre-determined clusters (assume k clusters). The main idea is to search for k centers from the given data set, one for each cluster. These centers should be placed in a cunning way because of different location causes different result. So, the optimal choice is to place the centers as much as possible far away from each other. The next step is to take each point belonging to a given data set and associate it to the nearest center. When no point is pending, the first step is completed and an early group age is done. At this point there is a need to re-calculate k new centroids as barycenter of the clusters resulting from the previous step.

At this point there is a need to re-calculate k new centroids as barycenter of the clusters resulting from the previous step. After obtaining these k new centroids, a new binding has to be done between the same data set points and the nearest new center. A loop has been generated. As a result of this loop, the k centers change their location step by step until no more changes could be done or in other words centers do not move any more. In this way, the given data set is divided into k stable sub-clusters. Mathematically, given a set of data points (x1, x2, …, xn), where each data point is a d-dimensional real vector, K-means clustering aims to partition the n data points into k (≤ n) clusters C = {C1, C2, …, Ck} so as to minimize the within-cluster sum of squares (sum of distance functions of each point in the cluster to the K center).

Hence, the algorithm aims at minimizing an objective function known as squared error function,(1)2𝑘𝑘 𝑗𝑗=1𝑥𝑥𝑖𝑖∈𝐶𝐶𝑗𝑗 where x i is the ith data point of the jth cluster, Cj; µj is the mean of jth cluster. The main steps of the direct K- means algorithm can be summarized as:�𝑥𝑥𝑖𝑖 − 𝜇𝜇𝑗𝑗�𝐽𝐽(𝐶𝐶) = ∑∑(1). Select an initial partition with K clusters; (2). Generate a new partition by assigning each pattern to its closest cluster center; (3). Compute new cluster centers; (4). Repeat steps 2 and 3 until cluster membership stabilizes.With this algorithm, all of the selected facets of an STL model can be clustered when using their facet normal vectors as given data points. For example, in Figure 2, the facets on the pair of parallel planar surfaces will be put into a same cluster and some of the facets on the pair of coaxial cylindrical surfaces will also be grouped into another same facet cluster due to similar facet normal vectors.

This will cause that disconnected facets will be clustered into a same cluster. Hence, when applying the K-means algorithm, there is no need to do feature recognition but just analyze the characteristics of facet normal vectors. Certainly, by analyzing the adjacent relationship and curvature of clustered facets, traditional shape features can also be formed and recognized by separating facet clusters or combining clusters that have connected facets. But as discussed above, for the layered manufacturing, it is better to consider facet groups with similar normal directions since they have similar production results. And when using these clusters to generate alternative build orientations, it is more efficient for the reduced duplicates. For example, in Figure 2, if applying traditional feature-recognition method to recognize the pair of cylindrical features, two planar surfaces features will be identified. The two features will generate similar alternative build orientations.

The two features will generate similar alternative build orientations. But, when using clustering method, the two features will be regarded as one cluster, which will generate a set of reduced alternative orientations.27Though the adopted clustering method is simple for understanding, fast and robust, the main disadvantage is that it requires apriori specification. This includes three user-specified parameters: number of clusters, k, cluster initialization (usually randomly choosing k of the sample data points), and distance metric for error measurement (usually using Euclidean distance). Among these parameters, the most critical choice is k [41]. In mathematics, this is usually a NP-hard problem. But in the application problem for discretely decomposing an STL model in this paper, it depends. There are two conditions: key surface areas of an STL model are already specified and no preference is given.

There are two conditions: key surface areas of an STL model are already specified and no preference is given. In the first condition, designers or process planning engineers already know which area of an STL model is critical. Hence, some of the facets of these key areas can be selected as cluster centers and the number of key areas is the number of clusters, k. In this way, the facets of key areas with other facets of other part of an STL model will be clustered together. The formed clusters can be used to generate alternative build orientations with associated generation rules. However, in the second condition, especially for STL models with complex freeform surfaces, designers or planning engineers usually don’t know or feel difficult to indicate the kea areas or give preference. There is a need to find an optimal number of clusters, k, to better decompose an STL model. As said above, in this case this is often NP-hard.

There is a need to find an optimal number of clusters, k, to better decompose an STL model. As said above, in this case this is often NP-hard. Hence, advanced computation algorithm is required.In this paper, to solve the problem of k-value specification, another statistical method, Elbow Method [42], is adopted. This method is an observation of the relationship between the value of the squared error function above and the number of clusters. Theoretically, if the k equals the number of given data points, the error would be zero, which means each point is a cluster. In this way, the clustering is meaningless. Hence, this method is dedicated to find the ‘elbow’ (optimal number of clusters), where the value of the squared error function drops the most dramatically and after that the value changes slowly. Figure 3 (a) shows the ‘elbow’, the circled point where k =2.

Figure 3 (a) shows the ‘elbow’, the circled point where k =2. However, for some data sets, there is no clear ‘elbow’ (Figure 3 (b)).Figure 3, typical relationships between the error function value and the number of clustersTo deal with this situation and obtain a better observation, a ‘gap statistic’ [43] method is used. This is an indirect observation to analyze the information contained in the formula given as below:, (2)∗ 𝐺𝐺𝐺𝐺𝐺𝐺𝑛𝑛(𝑘𝑘) = 𝐸𝐸𝑛𝑛{𝐽𝐽𝑘𝑘} − 𝑙𝑙𝑙𝑙𝑙𝑙(𝐽𝐽𝑘𝑘)The optimal number of clusters, k*, is the smallest k such that about the indirect observation method can be found in [43].. More details𝐺𝐺𝐺𝐺𝐺𝐺(𝑘𝑘) ≥ 𝐺𝐺𝐺𝐺𝐺𝐺(𝑘𝑘 + 1) − 𝑆𝑆𝑘𝑘+1By using the introduced two statistical methods above, a full or incomplete STL model can be decomposed in a discrete way to form both connected and disconnected facet clusters with or without apriori specification (specifying the key areas of an STL model) for generating meaningful alternative orientations.

In this paper, for simplicity, a central normal-based orientation generation rule is used. This means that the central point of each facet cluster is adopted as the normal vector of a build orientation for the facet cluster and then used as an alternative build orientation for the original full STL model. Actually, for specific application contexts to suite different characteristics of AM processes, different orientation generation rules can be set for facet clusters as did by traditional feature-based method in literature. Therefore, in general, the proposed statistical method can be described in three main steps:28(1). Determine the optimal number of facet clusters, k*. For this step, apriori specification or statisticalanalysis with Elbow Method can be used;(2). Decompose an STL model with the use of K-means clustering method; (3). Associate orientation generation rule for each obtained facet cluster.

Decompose an STL model with the use of K-means clustering method; (3). Associate orientation generation rule for each obtained facet cluster. In this paper, central normal vectorof each cluster is directly used as the direction of an alternative build orientation for the original STL model.The following section will present several examples to demonstrate the proposed discrete decompositionmethod and show how to use it to generate alternative build orientations.Demonstration ExampleIn this section, five STL models (Figure 4) are used to test the proposed statistical decomposition method for alternative build orientation generation. To be generic, an assumption is made that apriori specification is not available. Certainly, if preferences of designers or process planners are known, the clustering process would be much easier.

To be generic, an assumption is made that apriori specification is not available. Certainly, if preferences of designers or process planners are known, the clustering process would be much easier. But, in current practice, the preferences are usually hard to express or conveyed to process planner in the orientation determination work since some of them depend on the owners’ implicit knowledge. Assigning preferences would make the facet clustering process not stable. Hence, it is better to use a standard process, the proposed three steps above. At first, Part 1 in Figure 4 is used as an illustrative example to explain the full solving process. Then, for the left examples, only the obtained results are presented since the solving processes are the same.Figure 4, STL part models for demonstrationTo generate alternative build orientations for Part 1, the first step is to identify the optimal number of clusters of facets to be formed to decompose the STL model.

As proposed in this paper, the Elbow Method is adopted to analyze the ‘percentage of variance explained’ as a function of the number of clusters. The value of ‘percentage of variance explained’ is another way to represent the error function value as presented above. After computation with the input of all of the normal vectors of the facets of the STL file, the ‘elbow’, k*, optimal number of clusters, for Part 1 can be observed in Figure 5.29Figure 5, optimal number of facet clusters (the circled one) for Part 1Then, the next step is to applying K-means clustering algorithm to form facet clusters. Inputting the normal vectors of all of the facets and the optimal number of clusters, k*, the clustering result, including facet clusters and central normal vector of each cluster, for Part 1 can be obtained as outputs. As shown in Figure 6 below, 5 facet clusters are formed and colored.

As shown in Figure 6 below, 5 facet clusters are formed and colored. Table 1 presents the central normal vectors of facet clusters.Figure 6, facet clusters of Part 1 (each cluster is marked with one color)Table 1, central normal vectors of facet clustersCentral Normal Vector no1 no2 no3 no4 no5x 0 0 -0.5000 1.0000 -0.5000y 0 0 0.8660 0 -0.8660z 1.0000 -1.0000 0 0 0The last step is to use these obtained facet clusters to generate alternative build orientations for the STL model. As said before, for simplicity, only one generation rule is set for facet clusters in this paper. The rule is that the central normal vector of a facet cluster is set as the build orientation for benefiting the facet cluster and it is used as an alternative build orientation for the STL model. Hence, for Part 1 in the example, 5 alternative build orientations can be generated by using the five obtained central normal vectors. Figure 7 shows the generated alternative build orientations for Part 1.

Figure 7 shows the generated alternative build orientations for Part 1. Therefore, the first task of build orientation optimization problem in AM is solved. These alternative build orientations can be used for the later evaluation. Certainly, more orientation generation rules can be set for facet clusters according to specific application contexts. However, production knowledge is required for setting generation rules. Here, in this paper, the central normal30based generation rule is just used as a part of feasibility study for the proposed method. More generation rules and case studies will be investigated for some typical AM processes in the near future.Figure 7, alternative build orientations (arrow direction) for Part 1Similarly, for other part models, the generation procedures are the same.

Part 2 and Part 3 are decomposedinto 6 and 4 optimal clusters respectively (Figure 8).Figure 8, facet clusters of Part 2 and Part 3 (each cluster is marked with one color)As shown in Figure 8, a cylindrical and a conic surface features are decomposed into several facet clusters. This is different with traditional feature recognition methods which would treat the whole cylindrical and conic surfaces as a complete feature respectively. However, as said before, these connected facet clusters can also be combined into full traditional surface features by analyzing the adjacent relationship and curvature information. But, as proposed in this paper, there is no need to do this since these facet clusters will generate alternative build orientation set that will cover the set generated by using combined full surface features.

The main reason and logic is that if a build orientation which is beneficial to a full surface feature then it will also be beneficial to the segments of the full surface model. Though the proposed method can reduce large number of meaningless alternative build orientations, it also generates duplicates when encountering models with symmetric characteristic as shown by Part 2 and Part 3. Hence, if needed, the generated alternative build orientation set can also be refined. Perhaps, for simple part models, the efficiency of the proposed method is not so obvious. But for complex part models, this method will exhibit advantages as shown by the computation results of Part 4 and31Part 5. Figure 9 presents the facet clusters of Part 4. This part model is decomposed into 15 facet clusters.

Figure 9 presents the facet clusters of Part 4. This part model is decomposed into 15 facet clusters. Some clusters contain disconnected facets.Figure 9, facet clusters of Part 4 (each cluster is marked with one color)As shown in Figure 9, a set of parallel surfaces are clustered into a same facet cluster. Then, a comment central normal vector can be used to generate a comment build orientation which is beneficial for the facet cluster. However, if applying traditional feature recognition methods, these facet clusters which have parallel planar surfaces will be treated as many independent planar features and used to generate many duplicate build orientations. This will cause expensive computation cost for feature recognition, alternative orientation refining, etc. It can also be observed that the disconnected coaxial cylindrical surface segments of the part model are also grouped into comment facet clusters respectively.

It can also be observed that the disconnected coaxial cylindrical surface segments of the part model are also grouped into comment facet clusters respectively. Further, when facing part models with free form surfaces, process planners usually have difficulty to determine a set of finite alternative build orientations for evaluation since traditional methods (e.g. feature based methods) are not applicable since no obvious base plane or planar faces can be identified. But the proposed method can quickly generate optimal facet clusters to help process planners analyzing and generating a set of finite meaningful alternative build orientations. The computation result for Part 5 shows this advantage.32Figure 10, facet clusters of Part 5 (each cluster is marked with one color)As depicted in Figure 10, Part 5 with free form surface is decomposed into 31 facet clusters with disconnected facets.

Hence, 31 alternative build orientations can be generated for this part by using the central normal vectors of the facet clusters.Conclusion & Future WorkIn this paper, to solve the first task of build orientation optimization problem, alternative build orientation generation, with efficiency in AM, a facet clustering method is proposed. Using statistical clustering method, STL models can be decomposed into both connected and disconnected facet clusters that have similar production or design characteristics. These facet clusters can be used to generate alternative build orientations with associated generation rule. However, as a proof-of-concept study, this paper has not yet investigated the more orientation generation rules for real case study in concrete application context.

However, as a proof-of-concept study, this paper has not yet investigated the more orientation generation rules for real case study in concrete application context. Future work will be conducted to extract a set of practical orientation generation rule for some typical AM processes so as to make the proposed facet cluster based orientation generation method have industrialization chance. In addition, designers or process planners can also use the formed facet clusters to analyze the production quality distribution, e.g. surface roughness. For multi-axis AM systems, this method can help to determine the deposition sequence (orientation sequence), slicing, tool-path generation etc., which is the one of the main aims of the authors to develop the proposed method for supporting their following research work.References[1]. Wohlers, T. (2016). Wohlers report 2016. Wohlers Associates, Inc.[2]. Thompson, M. K., Moroni, G., Vaneker, T., Fadel, G., Campbell, R. I., Gibson, I.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFailure Detection of Fused Filament Fabrication via Deep LearningZhicheng Zhang* and Ismail Fidan¥ *Department of Mechanical Engineering ¥Department of Manufacturing and Engineering Technology College of Engineering Tennessee Technological University Cookeville, TN 38505AbstractAdditive Manufacturing (AM) is used in several fields and its utilization is growing sharply in almost every aspect of daily life. The focus of the current studies in the AM field is generally focused on the development of new technologies and materials. In addition, there is a limited number of research studies on the troubleshooting aspects of the AM processes. For the most commonly used Fused Filament Fabrication (FFF) process, the waste of material and time due to the printing errors are still an unsolved problem.

For the most commonly used Fused Filament Fabrication (FFF) process, the waste of material and time due to the printing errors are still an unsolved problem. The typical errors such as nozzle jamming and layer mis-alignment are inevitable during the printing process, and thus cause the failure of printing. It is a challenging task to clearly understand the physical behavior of FFF process with uncertainty, due to the phase transition and heterogeneity of the materials. Therefore, to detect the printing error, this research proposes a deep learning (DL) based printing failure detection technique. In this study, DL is utilized to monitor the printing process, and detect its failures. This newly developed DL framework was beta-tested with a commercially available FFF setup.

In this study, DL is utilized to monitor the printing process, and detect its failures. This newly developed DL framework was beta-tested with a commercially available FFF setup. The beta testing results showed that this technique could effectively detect printing failures with high accuracy.BackgroundAM is a set of technologies to produce three-dimensional (3D) objects from CAD (computer-aided design) models [1]. The term 3D Printing (3DP) is interchangeably used with AM [2]. AM has a broad range of applications in the fields of automotive, aerospace, medicine, etc [3, 4, 5, 6] because it has the following advantages:easy and fast production of complex workpieces, low cost and light weight end product,(cid:122)(cid:122) (cid:122) Environmentally friendly production, (cid:122) Less waste of material generated [7].

There are various methods of AM technology including FFF (Fused filament fabrication), SLS (Selective laser sintering), LOM (Laminated object manufacturing), STL (Stereolithography), etc [7]. However, the FFF technique is our particular interest since it is the most widely used one in industry and practical applications [8].2156Machine Learning (ML) is the process of using computer algorithm and statistical model to perform a specific task without using explicit instructions [9]. DL is one of the most popular ML methods, which has many applications in different areas of engineering [10, 11, 12]. DL is a class of ML algorithms using a cascade of multiple layers of nonlinear processing units for feature extraction and transformation [13]. In recent years, ML has absorbed many attentions in the area of AM to detect failure during printing and to optimize the process [14, 15, 16].Failure detection is one of the potential research problems in the field of AM [16].

It is worth investigation because it will help to eliminate waste material, shorten production time, and hence, save money for the users [16]. This is especially important in mass production since the solution of failure could help automating the process [17]. ML algorithm could easily be used for automatic detection of fabrication failures [16]. By using a DL method, parts which are built in an AM machine, could be classified as good (successful printing) or bad (failing printing) [18]. Then a camera could be used to detect and compare the produced part to the data set to predict whether the part is produced properly [19]. If the part is in good condition, the printing process continues, if not, the printing process pauses or stops. Failure detection of produced parts in AM is a critical task for broadening AM industrialization [20]. Widespread applications of this technology justifies the need of a method to reduce time, cost, and energy of the process.

Widespread applications of this technology justifies the need of a method to reduce time, cost, and energy of the process. An examination of current literature shows that a group used ML to do the failure detection [21], but there are no studies on using DL for failure detection parts. Thus, there is a need to do further research in this area.Structure of the Developed DL SystemThe aim of this study is to develop a DL algorithm to detect the part fabrication failures. The newly developed DL algorithm is applied to the 3D printer and the results are analyzed. Schematic of the system is shown in Figure 1. Firstly, the digital file of the part to be 3D printed is sliced in the slicer software, i.e. Cura. After slicing, the numerical g-code is generated and is sent to the 3D printer. The 3D printer runs these commands and starts extruding the material.

Cura. After slicing, the numerical g-code is generated and is sent to the 3D printer. The 3D printer runs these commands and starts extruding the material. The pictures of the fabrication process are taken by a GoPro camera, since the convenience and handiness, while the part is being built. Then, the images are sent to the computer where classifications into valid or invalid conditions of the process occur. The DL algorithm processes and analyzes the images based on the geometrical shapes. Overall, the developed DL system classifies the 3D printing as successful or failed after each layer.2157<~.,--_-_-_-_-_ -_ -_ ---l_J'J-t--_ -_ -_ -_ -_ -_ -_ -_ -....,Figure 1: Structure of the Developed SystemThe 3D printer used in this research is Ultimaker 3 [22]. The pictures of the layer-by-layer production are collected by a GoPro camera [23]. The GoPro is held by a separate frame and focused on the printing plane of the Ultimaker, and taking images during the printing process.

The GoPro is held by a separate frame and focused on the printing plane of the Ultimaker, and taking images during the printing process. In this research, the code is run in a Microsoft Surface Book laptop with the processor Intel(R) Core(TM) i7-8650U CPU @ 1.90 GHz 2.11GHz, and the installed RAM is 8.00 GB [24]. Moreover, the R language was used for the statistical computing and graphics.The flowchart of the code is shown in the Figure 2. There are five steps betweenstart and end.Input of images collected during 3D PrintingResize and reshape the imagesClassifica- lion and labeling of the imagesCreation of the model and doing he learningOutput of the accuracy and precisionFigure 2: Flowchart of the Code developed in this researchResultsBy changing the number of data, epochs and filters, the accuracy of the model (rate of correct predictions in all predictions) and processing time (time spent to do one single analysis) can be made to change.

In this research, all accuracy and processing time are calculated as the average of three different runs under the same conditions and the results are shown in Figure 3 and Figure 4.2158:~:~:;?' ~A 100g/100b e 80g/80b 60g/60b 40g/40bA I 00g/1 00b • 80g/80b 6 0g/6 0b 4 0g/4 0b25303540455020253035404550iurber of filters3.1mrnber of epochs.~·~ .~g -~ -~;--- ---:::::::::::::A 22 filters • 32 filters 42 filters 52 filtersAepochs 30 epochs " 4 0 epochs 50 e pochs8010012014016018020080100120140160180200nunberofdata3.3 Figure 3: Changes of Accuracymxnber of data(cid:122)(cid:122)(cid:122)(cid:122)In Figure 3.1, number of epoch is 30, this shows the change of accuracy on different number of filters. By increasing of the number of filters, the accuracy does not have a significant trend to increase. In Figure 3.2, number of filter is 32, this shows the change of accuracy on different number of epochs.

In Figure 3.2, number of filter is 32, this shows the change of accuracy on different number of epochs. By increasing of the number of epochs, the accuracy has a significant trend to increase. In Figure 3.3, number of epoch is 30, this shows the change of accuracy on different number of data. By increasing the number of data, the accuracy does not have a significant trend to increase. In Figure 3.4, number of filter is 32, this shows the change of accuracy on different number of data. By increasing the number of data, the accuracy does not have a significant trend to increase.Having accuracy is not enough to choose the most suitable parameters of the algorithm. In Figure 4, the change of processing time by different parameters is shown.21593.23.4(cid:122)(cid:122)calculation time vs. number of filters253035404550nlffberoffllters4.1calculation time vs. number of epochs20253035404550nt.mberolepochs4.2.calculation time vs.

number of filters253035404550nlffberoffllters4.1calculation time vs. number of epochs20253035404550nt.mberolepochs4.2.calculation time vs. number of data80100120140160180200size of data4.3 Figure 4: Changes of Processing TimeIn Figure 4.1, the graph shows the change of processing time based on different number of filters. In this research, the mean of processing time is 748.6 seconds and the standard deviation is 3.7 seconds. As the number of filters increases, the processing time remains constant. In Figure 4.2, the graph shows the change of processing time (T) based on different number of epochs (NE). The relationship between them is linear,2160T = 17.02 * NE + 107.68(1)(cid:122)As the number of epochs increases, the processing time increases. In Figure 4.3, the graph shows the change of processing time (T) based on different number of data (ND).

In Figure 4.3, the graph shows the change of processing time (T) based on different number of data (ND). The relationship between them is linear,T = 8.204 * ND - 206.643(2)As the total number of data increases, the processing time increases.The influence of each parameter on accuracy and processing time has been described, but it is still not enough to make the decision to choose the most suitable parameters.Evaluation of the AlgorithmIn the DL algorithm, accuracy is not the only character to judge how good the algorithm is. Besides the accuracy, precision is another important parameter to evaluate the algorithm.In this research, the images are classified into valid and invalid categories by the DL model. The predictions could be classified into four groups, which are defined as True Positive (TP), False Positive (FP), False Negative (FN) and True Negative (TN) [25]. •In TP case, the image reflects a valid print, the prediction of the algorithm is valid.

•In TP case, the image reflects a valid print, the prediction of the algorithm is valid. In FP case, the image reflects an invalid print, but the prediction of the algorithm is valid. In FN case, the image reflects a valid print, the prediction of the algorithm is invalid. In TN case, the image reflects an invalid print and the prediction of the algorithm is invalid.These four cases are shown in the Figure 5.False Positive (FP) Reality : Invalid DL model prediction: ValidFalse Negative (FN) Reality : Valid DL model prediction: InvalidFigure 5: Four cases of predictions2161Accuracy (A) is the ratio of all correct predictions to all predictions.A = (TP + TN) / (TP +FP + FN + TN) (3)Precision (P) is the ratio of all correct predictions to all valid predictions.P = TP / (TP + FP)(4)If the difference between accuracy and precision is small, the algorithm behaves similarly on both valid and invalid classifications.

But if the difference between accuracy and precision is significant, the algorithm will only predict based on one of two parameters; accuracy or precision.To evaluate the algorithm, a new parameter in defined by the research team. Reliability Factor (RF) is the difference between accuracy and precision divided by the accuracy, i.e., like percent difference. The equation of RF is as follows:RF = (A - P) / A(5)RF can range from positive to negative values. If RF is positive, A > P, the correctness of whole prediction is higher than valid prediction, the algorithm is more reliable in invalid predictions than valid predictions. If RF is negative, vice versa.ConclusionIn this study, the research team developed a DL algorithm to detect the failure during the FFF printing process. The number of data, epochs, and filters used in the algorithm was changed to test the accuracy, processing time, and RF.

The number of data, epochs, and filters used in the algorithm was changed to test the accuracy, processing time, and RF. In all cases of various parameter values, the algorithm was able to detect failure during the FFF printing process.By considering the accuracy, processing time, and RF, it is enough to choose the most suitable parameters of the code, which are 22 filters, 30 epochs and 120 images (60 good images and 60 bad images), because:(cid:122)(cid:122)(cid:122)This case has a high accuracy, 0.78, while the highest accuracy of all cases is 0.8. It could detect most failure during the FFF process. The number of epochs is 30, number of data is 120, which are all in the low level, so the processing time would not be long and the computational power would not be occupied a lot.

The number of epochs is 30, number of data is 120, which are all in the low level, so the processing time would not be long and the computational power would not be occupied a lot. The RF is 0, which means that this case behaves an equal prediction on good and bad images; avoiding the influence from false negative predictions.Thus, this algorithm could be used in the classification process. It could do a satisfactory prediction on both good and bad printing parts and would not occupy plenty of computational power. With the help of this DL algorithm, the failure during the 3DP process would be detected and the process would pause or stop in order to avoid the waste of materials generated by the bad printing process.- - -AcknowledgementsThe technical and editorial support provided by Mr. Mahmoud N Mahmoud, Mr. Astrit Imeri, Mr. Mahdi Mohammadizadeh, Dr. Yunbo Zhang, and Dr. Michael Allen is greatly appreciated.2162References[1] Gebhardt, Andreas.

[21] Delli, Ugandhar, and Shing Chang. "Automated process monitoring in 3D printing using supervised machine learning." Procedia Manufacturing 26 (2018): 865-870. (EN).pdf’, [22] ‘Ultimaker https://ultimaker.com/download/61355/Ultimaker%203%20manual%20%28EN%29.p df, accessed on June 9, 2019 [23] ‘UM_HEROPlus_ENG_REVA_WEB.pdf’, https://gopro.com/content/dam/help/heroplus/manuals/UM_HEROPlus_ENG_REVA _WEB.pdf, accessed on June 9, 2019 [24] ‘surface-book-user-guide-EN.pdf’, https://download.microsoft.com/download/7/B/1/7B10C82E-F520-4080-8516-5CF0D 803EEE0/surface-book-user-guide-EN.pdf, accessed on June 9, 2019 [25] Koo or ‘Accuracy, https://towardsdatascience.com/accuracy-precision-recall-or-f1-331fb37c5cb9, accessed on June 9, 20193manualPingShung’sPrecision,Recall2164F1?’,.

e Failure of Wire-Arc Additive Manufactured Aluminum Alloys With Porosities Under Loadings as Observed by In-Situ X-Ray Micro-Computed TomographyRunyu Zhang*, Yuxin Jiao*, Christopher Paniagua*, Yi Tian*, Hongbing Lu*, and Wei Li*Department of Mechanical Engineering, The University of Texas at Dallas, Richardson, TX 75080, USAAbstractWire-arc additive manufactured aluminum alloys (WAAM 4043 Aluminum) are widely used in many industries. Porosities are known to exist within the WAAM aluminum alloys, which greatly reduces the usability and reliability of such parts. In this study, WAAM aluminum alloy samples with porosities are manufactured using a Fronius (TPS 320i) MIG/MAG welding and ABB (IRB 140) robot system. The porosities generated inside the samples and the porosity evolution under the uniaxial tension are observed using in-situ X-ray micro-computed tomography (μCT).

The porosities generated inside the samples and the porosity evolution under the uniaxial tension are observed using in-situ X-ray micro-computed tomography (μCT). The μCT system with an integrated mechanical loading frame provides in-situ volumetric images of the specimens while loadings are applied. The porosity evolution of the WAAM aluminum samples and the propagation of the internal pores are assessed. This work provides direct experimental observations and evaluations of the influence of porosities on the mechanical behavior of WAAM aluminum alloys under loadings.Keywords: Wire-arc additive manufacturing (WAAM), porosity propagation, in-situ X-ray micro-computed tomography (μCT)IntroductionWire arc additive manufacturing (WAAM) is the process of combining the gas metal arc welding (GMAW) process with additive manufacturing (AM), that the metal wire is melted as feedstock using the high-power electric arc to form the geometry and the parts with designated printing path.

Compared to the power-based AM process, the WAAM process has a significantly lower cost, better applicability to reflective metals, and is also more environmental-friendly [1]. However, like any other AM process, processing parameters directly influence the printing process, including microstructure evolutions [2,3], molten pool formations [4,5], etc., thus determining the quality of the printed parts. Porosities are known to be the primary defects in the aluminum WAAM parts, due to significant solubility differences of hydrogen in aluminum liquid and solid [6,7]. Postprocessing is often applied to mitigate and reduce porosity formed in the parts [8,9], however, their mechanical properties are still often compromised [10]. In this study, single- track and multi-layer aluminum WAAM samples are printed with different processing parameters, and the porosity generation and evolution under uniaxial tension are studied using an in-situ X-ray micro-computed tomography (μCT) system.

The volumetric images from the reconstructed μCT provide direct experimental observations of the porosity size and distribution inside the WAAM prints from different processing parameters, and the porosity evolution under loading can be used to validate simulations in the future study.1456Materials and MethodsAll single-track and multi-layer samples are prepared using the WAAM system of the ABB robot arm and Fronius welder as shown in figure 1(a). The motion control platform for this WAAM process is a programable ABB robot arm that allows for six degrees of motion. This platform allows for accurate tool positioning when working in a small area. A Fronius TPS 320ti is used to set the welding parameters of the WAAM process, and the main factors consist of the current (A) and voltage applied, the wire feed rate (WFS) (mm/min), and the inert gas flow rate (CFH).

One feature of the Fronius TPS 320ti system is selecting “wall thickness”, a setting for traditional MIG welding, by changing which the system current, voltage, and WFR are automatically adjusted. To observe the porosities inside the WAAM samples and the porosity evolution during the uniaxial tension, a Nikon 225 kV X-ray μCT system with an integrated mechanical loading frame (Psylotech, Evanston, IL) as shown in figure 1(b) is used. The X-ray μCT system is able to image the samples with voxel size up to 3 μm/voxel while applying the load to reveal the interior of the samples under loading, and such in-situ X-ray μCT systems have been used in many applications in additive for characterizing materials properties [11–13], and real-time monitoring manufacturing [14,15].Fig.

1 Equipment used in this study (a) The WAAM system of ABB robot arm and Fronius welder, and (b) a Nikon in-situ X-ray μCT system with a loading frame.All WAAM samples are printed on aluminum substrates cut to 25.60 mm x 50.40 mm x 12.80 mm. Preparation of the substrates consists of deburring after cutting to fit in the vice grip and a cleaned before printing began. Cleaning was done with a steel brush, acetone, and IPA ultrasonication for 5 minutes after the substrates were dried before printing. The printing material is a commercial wire filler of Al 4043 manufactured by Forney. All multi and single-layer prints used the Fronius MIG pulse power source as their arc mode. The deposition strategy consists of hatching between layers to maintain uniform layering. All individual printing tacks of each sample kept consistent WFS/TS ratio, welding intensity, shielding gas flow rate (31.20 CFH) of pure argon, and working tip distance to the substrate at 10 mm.

All individual printing tacks of each sample kept consistent WFS/TS ratio, welding intensity, shielding gas flow rate (31.20 CFH) of pure argon, and working tip distance to the substrate at 10 mm. Different scanning patterns and WFS/TS ratios were conducted between samples to obtain uniform geometry prints. All samples were1457preheated to 45 – 55 C by a heat gun and measured by a FLIR E5-XT infrared camera. Once the substrate was brought up to temperature, a single track was deposited and then cleaned with a steel brush with a 90-second dwell time. The nozzle was moved two-thirds of the track width and deposited in the same direction, followed by the same cleaning and dwell time.

The same printing process is repeated for multi-layer prints.Table 1 Tested Fronius wall thickness presetting and the corresponding printing parametersParametersCurrent (A) Voltage (V) Power (W) Wire feeding speed (inch/minute)0.10 81 17.7 1433.7270Fronius wall thickness presetting (inch)0.16 124 19.6 2430.40.20 153 21.3 3258.94004800.50 241 25.2 6073.2750Table 2 Printing parameters of multi-layer WAAM aluminum samples1Wall thickness (inch)0.50Printing speed (mm/s)8Shielding gas flow rate (CFH) 31.2Line width movement (mm)Height movement (mm)Tracks / layer1Number of layers120.20831.21130.16831.21140.10831.21150.201531.2423660.20631.2632370.20631.2732380.20831.26324As there are many parameters to adjust and optimize in WAAM, and the time and cost investment would be significant if each parameter were explored, the Fronius intelligent welding process is utilized to simplify the printing parameters.

The primary parameter studied for the optimum printing is the printing power, and with different “wall thickness” parameters selected, four single-track WAAM samples are prepared as shown in table 2. Both single-layer prints and multiple-layer prints are prepared; only a wall thickness of 0.20 inches is studied with varying printing speeds and print geometries for both printing parameters, as shown in table 2 for multi- layer and single-layer prints. A few factors are considered when designing the printing parameters, such as the wire feeding speed (WFS) over the traveling/printing speed (TS), the WFS/TS ratio, and heat input for temperature control. It is recommended that a WFS/TS ratio of 5 to 25 should1458be maintained for an acceptable wall print [4].

It is recommended that a WFS/TS ratio of 5 to 25 should1458be maintained for an acceptable wall print [4]. For a stable deposition rate and suitable print geometries, the WFS/TS ratio of 13 to 25 is held, and a temperature range of 45 to 55 °C is maintained before printing another track [8].Result and DiscussionThe reconstructed μCT of the four single-track samples are presented in figure 2, including samples 1 to 4, with printing parameters shown in table 2. As the input power increases from setting the wall thickness value at 0.1 inch in figure 2(a), to 0.5 inch in figure 2(d), the number of pores is increasing and larger individual pores are observed. Also, underpower while printing samples 3 and 4 in table 2 leads to surface unevenness and insufficient wire melt, as figures 2(a) and (b) show; whereas the excessive power input causes sample collapse and over-melting of the substrate as figure 2(d) shows.

Therefore, from the preliminary study of printing single-track WAAM samples to obtain the optimized input power, the presetting “wall thickness” of 0.2 inch with a current of 153 A, a voltage of 21.3 V, and wire feeding speed of 480 inch/minute is selected for printing multi-layer samples.Fig. 2 The reconstructed X-ray μCT of four single-track WAAM samples showing the three cross sections and the 3D rendering, including (a) sample 4, (b) sample 3, (c) sample 2, and (d) sample 1, as shown in table 2.Four multi-layer WAAM samples are manufactured using the processing parameters listed as “sample 5” to “samples 8” in table 2, and then cut into a cylindrical1459shape with a length of 30 mm and a diameter of 10 mm using a wire electrical discharge machine (EDM). All four samples are imaged using the X-ray μCT system to inspect the porosity inside as shown in figure 3.

All four samples are imaged using the X-ray μCT system to inspect the porosity inside as shown in figure 3. Based on printing parameters listed in table 2, all four samples are manufactured with the same input power and shielding gas flow rate, whereas the printing speed, the number of tracks per layer, and the total number of layers vary. From μCT images in figure 3, there is significantly less amount of porosity inside sample 5 compared to the other three, which indicates that a travel speed of 15 mm/s is the optimum value in the study. Although sample 5 is printed with the largest number of tracks per layer and the most number of layers in total among all four samples, it is often believed that the travel speed is the most influential factor among all three parameters here in determining the quality of WAAM parts, due to its impact on the input power into the system, the molten pool formation, liquid solidification rate, etc. [1,4,6,7,15].1460Fig.

[1,4,6,7,15].1460Fig. 3 Cross-sectional view of the reconstructed X-ray μCT of four multi-layer WAAM samples, including (a) sample 5, (b) sample 6, (c) sample 7, and (d) sample 8, as shown in table 2.1461Fig. 4 The reconstructed μCT showing the sample and the porosity inside (a) at the intact state, (b) and (c) at the loaded state under uniaxial tension, and (d) the porosity evolution before and during the uniaxial tension loading.To observe the porosity evolution in the WAAM sample, the center portion of sample 5 is further lathed to a reduced section with a gauge length of 10 mm and a diameter of 6 mm, and then the lathed sample is loaded with a uniaxial tension on the in- situ X-ray μCT machine. The sample in the intact state and loaded state are imaged, and the reconstructed μCT is shown in figure 4.

The sample in the intact state and loaded state are imaged, and the reconstructed μCT is shown in figure 4. All porosities are segmented from the μCT using Otsu’s method based on the grayscale value, with a minimum pore diameter of 100 μm as the threshold, as pores with a diameter larger than 100 μm are often considered as the big and inhomogeneously distributed process pores [6,8], and of interest in this study. As figure 4(d) indicates, where the pores at intact and loaded states are marked in red and green, respectively, pores are elongated from a relatively uniform spherical shape into an ellipsoidal shape. And the volume fraction of those pores over the entire porosity inside the sample (excluding pores with diameters smaller than 28.6 μm) at the intact state and loaded state are determined to be 40.2% and 67.8%, respectively.

This value suggests that smaller pores form into bigger pores during the uniaxial tension, likely due to voids coalescence and cavity growth.ConclusionExperiments on printing WAAM samples with different processing parameters are conducted to explore the porosity generation primarily influenced by the input power and printing speed. And the porosity evolution under uniaxial tension is studied using an in- situ X-ray μCT system. It is found that upon tensile loading, the volume fraction of large pores with diameters larger than 100 μm over the entire porosity inside the sample increases dramatically from 40.2% to 67.8%, due to voids elongation, nucleation, and1462coalescence.

This study provides insight into the porosity formation and evolution inside the WAAM samples under the uniaxial loading, and is valuable for validating simulation models in future studies.AcknowledgmentsThe authors acknowledge support by the Department of Energy, National Nuclear Security Administration, Predictive Science Academic Alliance Program (PSAAP) under Award Number DE-NA0003962 and DE-NA-0003525. We also acknowledge the support of NSF CMMI-1661246, CMMI-1636306, and CMMI-1726435, and the support by the Louis A. Beecherl Jr. endowed chair.References[1] B. Wu, Z. Pan, D. Ding, D. Cuiuri, H. Li, J. Xu, J. Norrish, A review of the wire arcadditive manufacturing of metals: properties, defects and quality improvement, Journal of Manufacturing Processes. 35 (2018) 127–139. https://doi.org/10.1016/J.JMAPRO.2018.08.001.[2] M. Seifi, A. Salem, D. Satko, J. Shaffer, J.J.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFast and Simple Printing of Graded Auxetic StructuresM. A. Carton* and M. Ganter*Department of Mechanical Engineering, University of Washington - Seattle, Seattle, WA 98195-2600AbstractOne of the great promises of additive manufacturing is the ability to build parts with volumetrically graded parameters that would be difﬁcult or impossible with traditional manufacturing. This paper presents a method of procedural generation and unsupported fabrication of 2D objects patterned with functionally graded auxetic (negative Poisson’s ratio) cellular structures using commercially available FDM printers. Several types of two-dimensional auxetic pattern are fabricated. The resulting printed objects exhibit a graded response to load, deforming corresponding to local pat- terning.

Several types of two-dimensional auxetic pattern are fabricated. The resulting printed objects exhibit a graded response to load, deforming corresponding to local pat- terning. Deformation is studied using imaging of loaded structures and applications in several areas are considered.1IntroductionOne advantage of additive manufacturing over conventional manufacturing is the possibility of building in complex internal structure essentially for free. Among the design space that has re- cently become accessible are functionally graded objects with spatial variation in their material properties (such as stiffness, index of refraction, and electrical conductivity). These properties can be tailored by varying materials or by making a graded internal structure - a varying metamaterial - which causes the structure to have a different material property than its component material or materials.

These properties are common among plant and animal structural materials such as wood and spongy bone. A number of graded cellular or lattice structures have been investigated recently, for instance for in applications of energy absorption, piezoelectricity, and small lenses with vary- ing index of refraction. This paper presents a framework for procedural generation of a particular class of printable lattices: graded cellular auxetic parts – that is, materials with a varying negative Poissons ratio.When expanded in one dimension these materials expand in one or both orthogonal dimensions. Anisotropic auxetic materials can have a Poisson’s ratio of well under -1. With rare exceptions [1, 2]) this property is unique to artiﬁcial metamaterials. Some of the applications that have been investigated are sensors [3, 4], fasteners [5], textiles [6, 7, 8], protective equipment [9, 3, 10], and medical devices [3]. Auxetics are also being eyed for their property of sinclastic curvature (i.e.

Auxetics are also being eyed for their property of sinclastic curvature (i.e. positive Gaussian curvature in a sheet upon bending in one direction) which is desirable for the honeycomb composites commonly used in aerospace applications [11, 12, 13, 14].2270Figure 1: The re-entrant auxetic structures implemented in this project Left: Re-entrant octagon (bow structure). Center: Re-entrant star. Right: Double arrow.The properties of auxetic materials have been investigated at least since the 19th century [15] but due to their complex internal structures they have been limited in practice until the recent de- velopment of computer-aided design and manufacturing. Graded-auxetic and mixed auxetic and non-auxetic materials may be even more valuable.

Graded-auxetic and mixed auxetic and non-auxetic materials may be even more valuable. Non-auxetic materials with auxetic inclusions can have a higher Young’s modulus than each material alone [16] and can exhibit auxetic behavior as a whole or even negative thermal expansion [17] and can be designed for unusual load and bend- ing responses [18, 12, 19, 13]. Graded foam structures have shown promising results for improved impact resistance and energy absorption, providing superior response to uniform equivalents under some impact conditions [20, 21, 19]. Parts have been manufactured in a number of ways. Gluing re-entrant and convex cut-fold hexagonal honeycombs produces a sectioned object [22]; ﬁner gra- dations can be achieved with metal forming and welding [10], and much more easily with additive manufacturing [23, 24].This paper presents a framework for procedural generation of printable graded auxetic parts.

The patterns in ﬁgure 1 are generated parametrically, with their geometry determined by a user- deﬁned spatial function. The resulting models, generated in less than two minutes, are simple to design and print.2 MethodsThis section outlines the method used in this code base for procedurally modeling graded cel- lular structures. Models were generated in Grasshopper, the parametric scripting language of Rhinoceros 3D (RhinoBETA ver. 6.16.19169.14484 for macOS) using the package Clipper to gen- erate curve offsets [25]. This language allows procedural generation of NURBS objects. The code discussed in this section is available at https://github.com/mollyacarton/auxetic.The three patterns chosen for this code are all re-entrant auxetics. Re-entrant cells exhibit auxetic properties when their deformation is dominated by ﬂexure at the joints of the cell [26].

Re-entrant cells exhibit auxetic properties when their deformation is dominated by ﬂexure at the joints of the cell [26]. As the cell expands or contracts in one direction, the concave side or sides of the cell open, causing2271the cell as a whole to expand or contract. It is easy to show geometrically that altering the internal angle of these types of cellular auxetic patterns alters the resulting Poisson’s ratio (for example see the discussion in [11, 27])Each cell type was modeled as a parametric unit structure. In order to create varying cellular patterns, one of the deﬁning parameters of the shape must be controllable by a user-deﬁned func- tion. In this paper, each of the three patterns chosen has been modeled with a joint in the pattern chosen as the control point. The position of this control point and the points that follow it by sym- metry are assigned according to a mapping from the values of the spatial input function.

The position of this control point and the points that follow it by sym- metry are assigned according to a mapping from the values of the spatial input function. The rest of the shape of the cell is then determined according to this position.For example, in the star pattern seen in the center of ﬁgure 1, the control point is the endpoint of the straight segment connecting each star. Moving this point towards or away from the center of the cell controls the internal angles of the star shape and thus the deformation response of the pattern. When this point is moved away from the center of the cell the internal angles at the points reach perpendicular and the star becomes a square; the pattern is no longer re-entrant and therefore no longer auxetic. As the control point is shifted inward, the internal angle becomes more acute and the theoretical Poisson’s ratio increases in the negative direction (top in ﬁgure 2).

As the control point is shifted inward, the internal angle becomes more acute and the theoretical Poisson’s ratio increases in the negative direction (top in ﬁgure 2). The bow pattern (center in ﬁgure 2) similarly consists of octagons with two concave sides, with deformation controlled by the “waist” points of the cell shape. The arrow (sometimes called double arrowhead, bottom in ﬁgure 2) is controlled by the point of the arrow.The code functions as follows. Generate a square grid in the surface-local coordinates (a UV mesh). Then evaluate the spatial function at the center of each cell and map the output to the chosen domain of the controlling parameter. This determines the shape of the individual cell. For some patterns, like the bow and arrow shapes shown in 2 center and bottom, the cells share boundaries. Unlike the star pattern, it is necessary to choose which cell affects which part of the boundary in order to make the boundaries contiguous despite the changing cell shape.

Unlike the star pattern, it is necessary to choose which cell affects which part of the boundary in order to make the boundaries contiguous despite the changing cell shape. In the algorithm, this step is referred to as rectifying boundaries. This is deﬁned by the point that is chosen as a control point; as explained in the previous paragraph, the control points in the bow and arrow patterns both affect the neighboring cells as well as the cell in question. It is possible to classify patterns under “neighbor independent cells” and “neighbor dependent cells” as this affects the localization of response in the pattern.Once the cells have been generated, the lines are offset by the speciﬁed distance and a solid printable part extruded from the 2D pattern. Any additions to the model that are not part of the main cellular pattern, like solid bars for applying load to the ﬂexible part, can be added parametrically or by hand at this step.

Any additions to the model that are not part of the main cellular pattern, like solid bars for applying load to the ﬂexible part, can be added parametrically or by hand at this step. The parameters to fully deﬁne the part depend on the shape, but shared between all the patterns are printing volume parameters such as line width and extrusion depth, as well as grid scale and number of cells in the U and V directions. Treating the structure as a pattern on a UV mesh also provides a basis for mapping to a non-planar surface, although the parts that have been printed so far are planar.2272Figure 2: Linear gradation of the three patterns (from top: star, bow, and arrow), open on the left to closed on the right.

As the pattern becomes more closed and the internal angles more acute the magnitude of the Poisson’s ratio increases.2.1 PrintingThe graded structures shown in ﬁgures 3 and 5 were printed on a Creality Ender 3 commercial FFF printer, which has a removable ﬂexible magnetic bed, an advantage for removing well-adhered but delicate structures. The parts in 3 are PLA (polylactic acid) and the part in 5 the more elastic NinjaTek Cheetah TPU (thermoplastic polyurethane) from Fenner Drives, Inc. Prints were sliced in Ultimaker Cura 2.5.0 with a .2 mm layer height. The print temperature was 220°C and the bed temperature of 60°C. TPU was printed at 20 mm/second. Post-processing was limited to removing stray threads of material left by over-extrusion. Each sample part was generated using a 15 mm quadrilateral base cell size.2273Figure 3: Example printed parts in PLA.

Each sample part was generated using a 15 mm quadrilateral base cell size.2273Figure 3: Example printed parts in PLA. The cells are on a square grid with each unit cell 15 mm by 15 mm for all parts.22743 Case StudyThe example graded model and its print are shown in ﬁgures 4 and 5. This part is a strip of the arrow pattern, 4 by 14 cells, each cell 15 mm. The extrusion height is 5 mm and the line width is .45 mm, corresponding to one print bead width using a .4 mm nozzle. This model was generated in 11.8 seconds. The cells are graded with respect to the function f (x, y) = | sin(2πy)|, where the domain in y is [0, 1]. This gives a function with two sinusoidal peaks in y. The offset of the control point is between 20% and 80 % of the cell deﬁning scale of 15mm.The model was simulated in ANSYS Academic ver. 19.2 with a 5 N load applied to the upper edge. The printed part was then equivalently loaded.

19.2 with a 5 N load applied to the upper edge. The printed part was then equivalently loaded. Image measurements were performed in Δx/x0 Δy/y0 which is plotted in ﬁgure ImageJ [28], giving an approximate experimental Poisson’s ratio 6. This varied between -2.0984 and -0.4426 in an approximately sinusoidal manner.Figure 4: Arrow structure with spatial function | sin(2πy)|. Left: Original undeformed model. Right: Deformation under 5 N load in the y direction applied to the top edge (ANSYS ver. 19.2 Academic). Figures not to scale.2275Figure 5: TPU print of the model from ﬁgure 4.

Right: Deformation under 5 N load in the y direction applied to the top edge (ANSYS ver. 19.2 Academic). Figures not to scale.2275Figure 5: TPU print of the model from ﬁgure 4. Left: Undeformed Right: Deformed.Experimental linear approximation of Poisson's ratio in sample graded part0.500.5l0 ~ -1C/l C 0~(/) ·5 -11.5 c..l22.532468101214Cell row in yFigure 6: Plot of approximate Poisson’s ratio of the TPU printed part shown in ﬁgure 522764 ConclusionsThis paper introduced a body of code that generates printable graded auxetic structures through parametric modeling. The code provided in the github repository is available openly for research purposes. An example demonstrates the power of these printed structures - it is possible to print an object that not only has a negative Poisson’s ratio but has a varying Poisson’s ratio, so that it expands by different amounts in different places.

This has possible applications in, for example, fasteners, strain guages, and piezoelectric devices. A part could be designed that contacts sensors in a speciﬁed way when exposed to a certain amount of loading, or causes a secondary activation by physical or electrical contact.The power of computational design and additive manufacturing is still being uncovered. With more freedom from manufacturing constraints, simple design rules can produce objects with com- plex behavior, and these resulting objects have only begun to be explored. There is still a need for efﬁcient approaches to procedural generation of lattices and cellular structures, including those with unique material properties like auxetics.References[1] J. Jiang and H. Park, “Negative poissons ratio in single-layer black phosphorus,” Naturecommunications, vol. 5, p. 4727, 2014.[2] A. Marmier, S. Biesheuvel, M. Elmalik, A. Kirke, M. Langhof, J. Paiva, J. Toudup, and K.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFast Prediction of Thermal History in Large-Scale Parts Fabricated Via a Laser Metal Deposition ProcessLei Yan1,4, Tan Pan1, Joseph W. Newkirk2, Frank Liou1, Eric E. Thomas3, Andrew H. Baker3, James B. Castle31Department of Mechanical and Aerospace Engineering 2Department of Materials Science and Engineering Missouri University of Science and Technology, Rolla, MO, USA 3The Boeing Company, St. Louis, MO 4Email: lyvh8@mst.edu, lei.aaron.yan@gmail.comAbstractLaser metal deposition (LMD) has become a popular choice for the fabrication of near-net shape complex parts. Plastic deformation and residual stresses are common phenomena that are generated from the intrinsic large thermal gradients and high cooling rates in the process.

Plastic deformation and residual stresses are common phenomena that are generated from the intrinsic large thermal gradients and high cooling rates in the process. Finite element analysis (FEA) is often used to predict the transient thermal cycle and optimize processing parameters; however, the process of predicting the thermal history in the LMD process with the FEA method is usually time-consuming, especially for large-scale parts. Herein, multiple 3D FEA models with simple assumptions on the heat source and its loading methods are compared and validated with experimental thermocouple data.IntroductionIn the laser metal deposition (LMD) process, a description of the transient temperature fields is important for subsequent analyses such as microstructure evolution, residual stress, and distortion. Many efforts have been put on the temperature field analysis and are mainly based on the finite element analysis (FEA) method [1-5].

Many efforts have been put on the temperature field analysis and are mainly based on the finite element analysis (FEA) method [1-5]. In literature, a commonly used method to mimic the material addition process is called the “dummy material method” [3] or the “element birth and death method” [5], which use active elements in the finite element model according to a predefined deposition strategy. The aforementioned methods usually include a computational intensive model as not only do the minimum required time increments in transient heat transfer analysis need to be met [6], but also a mesh convergence needs to be achieved, which requires a minimum element size relative to the laser beam diameter [5]. When applied, the conventional methods to predict transient temperature history in a large-scale part fabrication process increase the model size tremendously.

When applied, the conventional methods to predict transient temperature history in a large-scale part fabrication process increase the model size tremendously. In some cases, a precise description of the thermal history in the laser deposition process is not necessary and only a fast prediction of the peak temperature will be needed for the process plan. Herein, two methods called track-heat-source and layer-heat-source are proposed for fast prediction of transient thermal history in large-scale part fabrication via the LMD process.Experimental Set-up and MethodsA Ti-6Al-4V block was built on a DMG-Mori Lasertec 4300 3D hybrid machine with the following processing parameters: laser power 1680 W, laser beam spot size 4 mm, powder feed517rate 29.7 g/min, and travel speed 1000 mm/min. The experimental set-up is shown in Fig. 1. Fig. 1 (a) shows the thermocouple locations on the substrate. The deposit was fabricated according to the path plan shown in Fig.

The experimental set-up is shown in Fig. 1. Fig. 1 (a) shows the thermocouple locations on the substrate. The deposit was fabricated according to the path plan shown in Fig. 1 (b), where nine single-direction laser scans were applied and a perimeter pass was performed on every layer. To provide a flat deposition on all corner walls, the perimeter pass was doubled every other layer based on trial-and-error. Fig. 1 (c) shows the thermocouples were fixed on the substrate with high temperature thermally conductive paste and temperature was recorded with a data acquisition instrument. Two holes shown in Fig. 1 (c) were used for mounting the substrate on a base plate. The final product is shown in Fig. 1 (d). Due to the working condition differences among varies LMD systems, there is no comparison with other groups’ work and only focused on the experimental data obtained from our LMD system.

1 (d). Due to the working condition differences among varies LMD systems, there is no comparison with other groups’ work and only focused on the experimental data obtained from our LMD system. To quickly predict the transient thermal history of a large-scale part build on this machine, an FEA model suitable for this process was developed and the predicted temperature fields will be used for subsequent analyses such as oxidation, residual stress, and distortion in the future.518Fig. 1 Experimental set-up.

1 Experimental set-up. (a) thermocouple locations, (b) laser scanning pattern, (c) thermocouples on the substrate before deposition, (d) final deposit.Modeling ProceduresIn our previous work of simulating the LMD process, the heat source was modeled with a conventional method as thermal flux density (watt/m2) or heat generation rate (watt/m3) that moves place to place according to a scanning strategy [5,7], which requires a minimum element size (<1/6 laser beam diameter) to meet the requirement of mesh convergence assessment. This criteria results in a large number of elements and nodes in the meshed model and results in a long simulation time. To reduce the model size and computational time, and make it feasible for fast prediction of the temperature in the LMD process, especially for large-scale parts, some assumptions and simplifications were proposed and implemented in the model:1.

In the current model, two bolts were simplified as two blocks at two ends of the substrate, and the base plate underneath the substrate was also modeled as a block, as shown in Fig. 2 (a), where their material properties were assigned using the equivalent mass method as proposed in [5]; mass, thermal conductivity, and specific heat were kept the same, but density was modified according to the modeled fixture volume. As shown in Fig. 2 (a), the model was meshed with a non-uniform distributed element size and only used fine mesh at volumes of interest to reduce the model size and computational time. The final model contains total 14922 elements and 17150 nodes.The initial condition for the numerical simulation domain was set to a uniform temperaturefield equal to ambient temperature (300 K), which is defined as shown in (3).

As new elements changed from death to birth status, the outer surfaces related to boundary conditions were updated.,LaserIn this case, forced convection needs to be considered as reported in [10], where the decrease of h as the distance from the top edge of the wall increases was assumed to be linear as provided in (5):ℎ = −1.4 + 80,where  is the distance from the top edge of the wall to the point of interest. The convective heat transfer coefficient distribution at the last step is shown in Fig. 2 (b), where the substrate and fixture were set to free convection when  was beyond the deposit height and the convective heat transfer coefficient was equal to 30 W/m2/°C.520(2)(3)(5)521Fig. 2 FEA model set-up.

2 FEA model set-up. (a) meshed mode with the simplified fixture (red), the substrate (white) and deposit (white), (b) convection coefficient distribution, (c) temperature distribution at 11th layer with track heat source, (d) temperature distribution at 11th layer with layer heat source.The finite element model uses a fixed mesh size for all the substrates, deposits, and fixtures during simulation. To simulate the addition of material in the laser deposition process, a method called element birth and death in ANSYS was employed. The element death function was used to deactivate an element by setting its stiffness to zero, while the element birth function was used to activate an element by setting the right stiffness value. At the start of the simulation, the substrate elements were all activated. However, the deposit elements were activated sequentially to simulate material addition.

At the start of the simulation, the substrate elements were all activated. However, the deposit elements were activated sequentially to simulate material addition. The element birth and death method application in track-heat-source and layer- heat-source are demonstrated in Fig. 2 (c) and 2 (d), respectively. Fig. 2 (c) shows the tracked heat source process where the heat flux was applied to the last track of the 11th layer, followed by the volume heat generation rate. The perimeter scan was simulated after the last track was added, following the same heat source application sequence as the last track. In Fig. 2 (d), the layered heat source process is shown, with the only difference from the track-heat-source method being the material was added layer by layer instead of track by track.Results and DiscussionsThe predicted transient thermal history from the FEA model was validated with the thermocouple data and the comparison results are shown in Fig. 3.

3. TC1 was caught in the laser path during deposition and its measured data was peaked out. Only TC0, TC2, and TC3 measured522data were used in the model validation process. The comparison between TC0, TC2, and TC3 is plotted in Fig. 3 (a), which shows some variance from 0 to approximately 150 seconds and aligns well after that until the end of deposition. The mismatch around 150 seconds is attributed to the difference in loading method in simulation and reality. In simulation, the heat source was applied track by track, but in reality the heat source (laser beam) started from one end and moved to the other end of each track, which led to the locations where the thermocouples were attached to be heated up gradually and reached a peak temperature every time the laser beam passed over. After certain layers, the substrate was heated up and the temperature difference caused by the loading methods was reduced.

After certain layers, the substrate was heated up and the temperature difference caused by the loading methods was reduced. Heat input and heat loss at evaluated locations on the substrate attained equilibrium for a period of time. As the simulation progressed, equilibrium was broken as heat loss was larger than the heat input, which resulted in a temperature decrease. This phenomenon has been observed in both experimental and simulation data. Based on this validated model, the temperature history of the first track’s middle point from the first, ninth, and 18th layers were predicted and plotted in Fig. 3 (b), which shows the peak temperature is around 3600 K and close to the evaporation temperature of titanium. The layer heat source method validation results are shown in Fig. 3 (c), where the same trend is captured after about 150 seconds, meaning the simulation and experimental data are well matched.

The layer heat source method validation results are shown in Fig. 3 (c), where the same trend is captured after about 150 seconds, meaning the simulation and experimental data are well matched. The temperature difference at about 150 seconds also comes from the difference in simulation and experimental loading methods. The predicted middle point peak temperature from the first track of the first, ninth, and 18th layers is close to 4200 K using the layer heat source model and exceeds the evaporation temperature of titanium. The peak temperature overestimation comes from the lack of cooling time between each track or layer compared to the real experiments.Comparing these two methods, the track-heat-source method has a better resolution than the layer-heat-source method, where the former can capture more thermal details better than the latter method.

However, compared to the conventional method which applies the heat source spot by spot, both methods have a lower resolution.Considering the computational time, the track-heat-source method took about 55 minutes and the layer-heat-source method took about 5 minutes on a computer with an AMD Phenom™ II X6 1090T Processor 3.2 GHz and 16.0 GB RAM hardware. The rougher the prediction model implemented, the less computational time is needed and sacrifices accuracy of the results.523Fig. 3 Model validation results. (a) temperature comparison with track heat source method, (b) prediction of temperature at first, ninth, and 18th layers with track heat source method, (c) temperature comparison with layer heat source method, (d) prediction of temperature at first, ninth, and 18th layers with layer heat source method.524ConclusionsTo conclude, the current work extends the modeling method for fast temperature prediction in large-scale part fabrication using the LMD process.

The following are the salient conclusions in the paper:1. Track- and layer-heat-source methods have the capability for fast temperature prediction in the LMD process.2. Deposit peak temperature was overestimated in both methods and came from sacrificing modeling details for the computational time-saving purpose.3. The track-heat-source method has a higher resolution than the layer-heat-source method, but is relatively slower, taking about 11 times longer computational time.AcknowledgmentsThis project was supported by The Boeing Company through the Center for Aerospace Manufacturing Technologies (CAMT), National Science Foundation Grants #CMMI-1547042 and CMMI-1625736, and the Intelligent Systems Center (ISC) at Missouri S&T. Their financial support is greatly appreciated.References[1] Costa, L., et al.

Recommended values of thermophysical properties for selected commercial alloys. Woodhead Publishing, 2002.[10] Heigel, J. C., P. Michaleris, and E. W. Reutzel. "Thermo-mechanical model development and validation of directed energy deposition additive manufacturing of Ti–6Al–4V." Additive manufacturing 5 (2015): 9-19.526.

st Production of Technical Prototypes Using Direct Laser Sintering of Metals and Foundry SandChristian Wilkening EOSGmbH1. Introduction Currently most RP-Parts are used for visualisation, assembly checks and some functional testing. The next stage between these functional prototypes and the pre-series is technical prototypes, which are used for final functional testing and optimisation of the production process. These must not only have the same material as the series parts, but must also be manufactured with the same production process. Due to the second of these requirements layer manufacturing processes cannot be used to build these prototypes directly, but they can be used to make the negative moulds or tools.Two new RP-processes based on laser sintering now have the capability to produce the moulds and tools via very short and fast process chains.

Both have already been in commercial use in Europe for about one year.The first process manufactures tools for injection moulding of plastic parts by Direct Metal Laser Sintering (DMLS). The second process, called the Direct Croning Process (DCP), is used to build sand moulds and cores for sand casting of metal parts directly from Croning-Sand without any tools. These technologies have been developed by EOS GmbH, Munich, and are marketed under the names EOSINT M and EOSINT S respectively.2. Direct Metal Laser Sintering (DMLS) The most important application for laser sintering of metals is the production of tooling and moulds for injection moulded parts. Two fundamental approaches are possible: • • Direct liquid phase sintering (DMLS). EOS has concentrated on DMLS, because the indirect process chain makes it very difficult to achieve the narrow tolerances required in toolmaking without repeated iterations.

EOS has concentrated on DMLS, because the indirect process chain makes it very difficult to achieve the narrow tolerances required in toolmaking without repeated iterations. With DMLS it is possible to get very close to the final geometrical and material properties already in the thereby avoiding the loss of time and accuracy associated with direct sintering process, secondary sintering processes.Indirect laser sintering with organic binder followed by solid state sintering in a furnace /1/2.1. Material and process The basic composition of the DMLS material was developed and patented in the late 1980s by Electrolux Rapid Development, Finland for tooling production by pressureless furnace sintering. EOS now owns the exclusive worldwide licence to these patents. The composition of the bronze-nickel based powder was described at the last two SFF symposia in Austin /2/, /3/.

EOS now owns the exclusive worldwide licence to these patents. The composition of the bronze-nickel based powder was described at the last two SFF symposia in Austin /2/, /3/. By using the right parameters the typical shrinkage due to the liquid phase sintering can be completely compensated by an expansion of the material caused by diffusion of the133components so that this material has no net volume change during the laser sintering process. This eliminates the need for a high temperature environment, which is otherwise required to reduce the internal stresses induced during laser sintering of other metal powders. The DMLS process works at ambient temperature. Due to the small laser focus of 350 /lm and the comparatively high scanning speed of 300-800 mm/s the duration of heating is short enough to avoid oxidation even without inert gas atmosphere.

Due to the small laser focus of 350 /lm and the comparatively high scanning speed of 300-800 mm/s the duration of heating is short enough to avoid oxidation even without inert gas atmosphere. The phosphor content of the third material component inhibits oxidation of the powder mixture and also improves the wetting of the solid particles in the molten phase. This short reaction time of few milliseconds was the reason for the relatively poor mechanical properties of the first laser sintered samples of the initial powder mixture, which had been optimised for furnace sintering. The wetting and the diffusion reaction were not completely finished because the available time was simply to short.

The wetting and the diffusion reaction were not completely finished because the available time was simply to short. During the last nine months the mechanical properties and the reliability of the process have been significantly improved by adapting the material composition to the specific requirements of the laser sintering process.The most important variations are an icreased content of the liquid phase, an enlarged surface area for the diffusion and an optimisation of the laser and scanning parameters. The tensile strength of the green part has been improved from 81 N/mm 2 last year to currently 150 N/mm 2 and the bending strength from 150 to 300 N/mm2 The resulting porosity of these parts when sintered without shrinkage to obtain the highest accuracy is approximately 25%. The green parts can be machined just like an aluminium part by milling, grinding, drilling etc. and the strength is already sufficient for special applications like blow-moulding or vulcanisation moulds.

and the strength is already sufficient for special applications like blow-moulding or vulcanisation moulds. These figures can be improved by an infiltration process that fills the pores. The preferred infiltration fluid is a high temperature epoxy resin, because this secondary process has no influence on the accuracy due to the relatively low heat impact to the metal. The result of the infiltration is an improvement of the bending strength to approx. 400 N/mm2 and a smoothing of the surface.The hardness after the infiltration is 108 HB. The infiltration is driven by the capillary forces, only the rear side of the mould must be dipped into the resin. This procedure needs about half an hour, the post curing in a simple oven at 160°C needs about two hours. Preheating of mould and resin to 60°C accelerates this process and reduces the residual porosity. Silver-based infiltrants are currently in the evaluation phase.

Preheating of mould and resin to 60°C accelerates this process and reduces the residual porosity. Silver-based infiltrants are currently in the evaluation phase. Although the mechanical properties can be improved, the cost of the necessary equipment and the loss of accuracy due to a shrinkage of more than 2% are a major disadvantage 13/.The negligible shrinkage of the DMLS process enables an accuracy of epoxy infiltrated parts of 0.05 to 0.1 % in the first trial, which is even comparable to stereolithography. The limit for the accuracy is the mean particle size of 35-40 /lm. Typical layer thicknesses of 80-100 /lm are used to obtain a comparable resolution in the z-direction. The roughness of the infiltrated part is about Ra = 3.. 5 /lm, which is comparable to a rough EDM surface. With standard manual polishing, a roughness of Ra < 1 /lm is possible. The heat conductivity of the sintered tools of 110 WImK is slightly lower than the value of aluminium.1342.2.

With standard manual polishing, a roughness of Ra < 1 /lm is possible. The heat conductivity of the sintered tools of 110 WImK is slightly lower than the value of aluminium.1342.2. System All EOS laser sintering machines are designed according to the specific requirements of the used material and optimized for the intended application. The main difference between the sintering systems for metal and for polymers is the design of the optics and the laser, because the much higher sintering temperature of approximately 900 °C requires a higher power density. Therefore the focus diameter was reduced to 350 J-lm and the laser power was increased to 250 Watts. The power density thereby was increased from 25 W/mm2 for plastic powders to 700 W/mm2 for metal. More powerful lasers are still in the test phase. Since no preheating of the powder is needed, complicated. The working area of the EOSINT M 250 machine is 250 mm square, the maximum part height is 150 lTIIll.

Since no preheating of the powder is needed, complicated. The working area of the EOSINT M 250 machine is 250 mm square, the maximum part height is 150 lTIIll. With a density of the sintered material of 6, 1 g/cm3 the maximum weight of the part is more than 50 kg.the design of the process chamber is less2.3. Applications The most important application of DMLS is the production of prototype tooling for injection moulding. The process is especially useful for complex moulds. If the geometry of the tool is simple enough to be manufactured just by milling, the rapid tooling processes generally cannot compete against high speed milling machines. However if EDM is required for some details in the tools, the DMLS process is usually faster and often less expensive than the conventional methods. Sintering and infiltration of the mould inserts needs only 1-2 days, depending on the size.

Sintering and infiltration of the mould inserts needs only 1-2 days, depending on the size. The time needed for CAD design of the tool and for finishing and mould assembly is the same as in the conventional process for aluminium tools. The building accuracy of the split planes is sufficient that only manual finishing is needed, thereby eliminating the need for the expensive and time consuming design of an offset on these planes and an additional machining step. Using DMLS several service bureaux have made many projects in a total processing time of only one week, including tool concept, CAD design of the tool, DMLS, infiltration, finishing, mould assembly and injection moulding.Fh~ 1: Top and bottom view ofa motor holder, PA66/35 % GF, dia 230mm135ambient temperature this coating is dry and hard. The total resin content of the mixture is about 5 %, which is comparable to standard foundry sands.

The total resin content of the mixture is about 5 %, which is comparable to standard foundry sands. The use of this material for building moulds and cores by laser-sintering has been patented by EOS. At a temperature of 150 - 200°C the polymerisation reaction starts and leads to an irreversible thermoset bonding between the sand particles. In the conventional foundry process the Croning sand is poured over hot steel tools having the positive part geometry, in the Direct Croning Process the energy input is completely provided by a CO2 laser. No additional preheating of the process chamber or the sand is necessary. The process also does not require an inert gas atmosphere. The sand costs are about 1% of typical stereolithography resins and unsintered sand can be reused several times. is neglegible « 0.1 %).

The process also does not require an inert gas atmosphere. The sand costs are about 1% of typical stereolithography resins and unsintered sand can be reused several times. is neglegible « 0.1 %). The Due to the low binder content the shrinkage of the material accuracy of the sintered moulds is therefore only limited by the positioning accuracy of the scanners and the particle size (for fine details).The surface quality of the cast parts is comparable to the typical figures for conventional sand casting. If a smoother surface is required, these specific regions can be coated with the commonly used ceramic blackwash.DCP without blackwash DCP with blackwash Green sand Croning shell50 - 80 Jlm 25 - 30 Jlm 50 - 70 Jlm 20 - 40 JlmTable 1: Roughness Ra of aluminium castings depending on the mold materialGenerally every alloy that can be used for sand casting can also be used with DCP moulds.

So far various alloys of aluminium, magnesium, iron and steel have been cast sucessfully.3.2. System The EOSINT S700 sand laser-sintering machine is the first RP system to use a twin-laser setup. Each laser has its own scanner and they operate simultaneously and in parallel to effectively double the building speed. Each of these scanner heads has a working area of 380mm square, which allowing for a small overlap in the middle gives a total building area of 720 mm by 380 mID. The maximum part height is also 380 mID. The sintered parts are removed from the machine in a container and separated from the loose sand in an external recycling station. The machine can immediately start the next job. The high maximum part weight of more than 200 kg requires a special handling system for the part containers, which is supplied with the machine.

The machine can immediately start the next job. The high maximum part weight of more than 200 kg requires a special handling system for the part containers, which is supplied with the machine. A pneumatic handling system is also included in the machine for automatically refilling with sand without interrupting the building process.136The twin head design and the high scanning speed of 2000 mm/s enable a very high process productivity. Depending on the complexity of the geometry the typical rate is 0,8 - 1.5 kg part weight per hour.3.3. Applications Like laser sintering of thermoplastic materials, DCP has very few limitations on the part complexity. Almost every geometry which can be cast can also be built with this process. The limits of sand casting (e.g. a minimum wall thickness of 1.5-2 mm) can be achieved without problem.

Almost every geometry which can be cast can also be built with this process. The limits of sand casting (e.g. a minimum wall thickness of 1.5-2 mm) can be achieved without problem. DCP often allows the design of the mould to be significantly simplified, for example no draft angles are needed, undercuts are easily possible and the number of the necessary core-pieces can often be reduced. (see Fig 3)Fig 3: Crankshaft with 9 hollow bearingseats, casted in a two-part mouldAlthough DCP has many applications including hydraulic parts, pump components etc., the most important one is the production of prototype parts for engine development. The potential of this process is impressively demonstrated by the example of cylinder head development. Due to the extremely complex geometry of this component with internal water cooling channels, valves and many inlet/outlet channels, prototype tooling is very complicated and difficult to manufacture.

The time required between CAD design and the first casting using conventional methods is 15-20 weeks. Because the design of the cylinder head has a very significant influence on the engine performance, many iterations and optimisations are necessary for this component. Therefore this part is usually the most time-critical component which determines the duration of the engine development. The build time for a I-cylinder head using DCP is only 2 days, for a 4 cylinder head about one week. Together with the additional design of the mould and the runner/gating system which is needed anyway, the total processing time of a new head can be reduced to 2 - 3 weeks, which means a time saving of 80%! The DCP-produced cylinder heads are virtually identical to series parts and can run in test engines.137These figures are based on the experiences of the beta customer BMW, who was involved in the development of the system and the process from the start.

With the two units that were installed in their foundry 6 and 12 months ago they have already cast almost 100 different engine designs. The other partners of the beta test programme are Mercedes Benz and several service bureaux that work for the automotive industry.Fig 4: Core for the water cooling jacket of a I-cylinder test engineThe high build speed makes it possible to produce not only single prototypes but also small series. The biggest order so far for one of the service bureaux was the production of 1000 complicated cores (120 x 120 x 80 mm) for a water jet system.4. Conclusions The production of technical prototypes (i.e. in series materials and with series production techniques) by Rapid Prototyping was until recently only possible via Rapid Tooling process chains involving at least one conversion process, due to the limitations of the available RP materials.

The recent introduction of Direct Metal Laser Sintering (DMLS) and the Direct Craning Process (DCP) has for the first time enabled the direct production of tooling and moulds for injection moulding and sand casting respectively, and thereby the fast and efficient production of technical prototypes in both plastics and metals. This marks a very significant advance in Rapid Prototyping and Tooling technology. The ability of DCP to directly manufacture one-off and small-series parts with identical quality to conventional sand-cast series parts has also for the first time made the breakthrough into true Rapid Manufacturing.138Since neither the sharpness of the edges nor the levels of form and dimensional accuracy achieved meet the requirements of a pre-series tool, the contour of the part is milled. The upper surfaces also undergo a milling operation to keep the distance between the upper edge of the last deposited layer and the lower edge of the powder nozzle constant, Figure 2.

The suitability of this method for automation in comparison to e.g. conventional 5-axis milling is a powerful advantage. By virtue of the layer-by-layer nature of this technique, the CAD data may be processed quickly and with considerably less effort.the Fraunhofer IPT is essentially a high speed milling machine supplied by the Albrecht Roders company on which a laser generating head has been mounted, Figure 3. To date, simple geometries have been built using different materials in order to investigate the principle underlying the CMB process (Figure 3).The test facility atSelective Laser SinteringCurrent approaches to sinter metallic powders are concentrated on the direct or indirect sintering of the chosen material. Indirect sintering, deploying a binder phase (e.g. a polymer), has the advantage that only the binder material needs to be melted, most of the time at a low temperature allowing manufacturing of metallic parts using conventional plastic sintering machines.

However, the workpiece must be debindered and infiltrated, causing loss of accuracy and prolonging the time needed for manufacture. Direct sintering is aimed at melting the chosen material during the SLS process directly, eliminating the need for debindering. However, since processes currently industrially available leave the parts with a porosity of about 70%, infiltration is still necessary to achieve full density. Another problem related to direct sintering of metallic139These technologies have also introduced new technical innovations which are pushing RP technology forwards. EOSINT M 250 is the first commercial system to build parts directly in metal, while EOSINT S 700 is the first RP system worldwide to use two laser-scanner systems in parallel for increased productivity.Due to the considerable time-saving benefits of these new technologies they have established themselves very rapidly in Europe.

EOSINT M and EOSINT S systems have been in use at service bureaux and industrial end-users since the summer of 1995, including such reference names as BMW, Merrcedes and Electrolux. These companies are already experiencing the technical and commercial advantages of true Rapid Tooling, and the growing user-experience being generated by the rapidly expanding customer base is indicating a number of exciting new applications which will be reported in future.References/1/Badrinarayan B. and Barlow, 1. W.,"Metal Parts from Selective Laser Sintering of Metal-Polymer Powders", Solid Freeform Fabrication Symposium, Austin 1992, 141 146/2/M. Sindel, T. Pintat, M. Greul, O. Nyrhila, C. Wilkening,"Direct Laser Sintering of Metals and Metal Melt Infiltration for Near Net Shape Fabrication of Components", Solid Freeform Fabrication Symposium, Austin 1994, 94-101/3/T. Pintat, M. Greul, M. Greulich, C.

Pintat, M. Greul, M. Greulich, C. Wilkening,"Accuracy and Mechanical Behavior of Metal Parts Produced by Laser Sintering", Solid Freeform Fabrication Symposium, Austin 1995, 72-78140.

st Recoating Methods for the Projection-based Stereolithography Process in Micro- and Macro-ScalesYayue Pan, Yong Chen*, Chi Zhou Daniel J. Epstein Department of Industrial and Systems Engineering University of Southern California Los Angeles, CA 90089, U.S.A.Author of correspondence, Phone: (213) 740-7829, Fax: (213) 740-1120, Email: yongchen@usc.eduABSTRACTThe purpose of this paper is to present a recoating method for the development of a direct digital manufacturing (DDM) process that can be an order of magnitude faster than other currently available DDM processes. In the mask-image-projection-based Stereolithography (MIP-SL) process, projection light controlled by a Digital Micromirror Device (DMD) can quickly cure liquid photopolymer resin in a whole area; a fast recoating method is required for achieving truly high-speed fabrication. We investigate the bottom-up projection system in the MIP-SL process.

We investigate the bottom-up projection system in the MIP-SL process. For the macro-scale MIP-SL process, a two-way linear motion approach has been developed for the quick spreading of liquid resin into uniform thin layers. In comparison, a direct pull-up motion can be used in the micro-scale MIP-SL process. The system design and related settings for achieving a fabrication speed of a few seconds per layer are presented. Additionally, the hardware, software, and material setups for fabricating three-dimensional (3D) digital models are presented.

Additionally, the hardware, software, and material setups for fabricating three-dimensional (3D) digital models are presented. Experimental studies using the developed testbed have been performed to verify the effectiveness and efficiency of the presented fast MIP-SL process.KEYWORDSAdditive manufacturing, high-speed fabrication, fast recoating, stereolithography.1 INTRODUCTIONLayer-based Additive Manufacturing (AM) processes such as Stereolithography Apparatus (SLA) can fabricate parts directly from Computer-aided Design (CAD) models without part-specific tooling or fixtures. As a direct manufacturing approach, AM processes can cost-effectively fabricate truly complex three-dimensional (3D) shapes that were previously impossible. Unlike traditional prototyping approaches that take days, AM-based rapid prototyping can build physical objects in hours.

Unlike traditional prototyping approaches that take days, AM-based rapid prototyping can build physical objects in hours. Due to such time and cost benefits, AM processes have been widely adopted in the product development process for building prototypes of a design.Although the speed of AM systems has significantly increased over the years, the building process of a moderate sized 3D model is typically measured in hours. In a recent NSF workshop on developing the roadmap for AM [1], the development of AM machines with higher throughput was identified to be critical for future rapid manufacturing requirements. Future high-speed AM systems require new approaches, evolving from point-processing or line-processing methods such as a laser or an extruding nozzle, to area-processing or volume-processing methods.In this research, we investigate the building speed of an area-processing approach that is based on a Digital Micromirror Device (DMD).

A DMD is a microelectromechanical system (MEMS) device that enables one to simultaneously control ~1 million small mirrors that turn on or off a pixel at over 5 KHz. Using this technology, a light projection device can project a dynamically defined mask image onto a resin surface to selectively cure liquid resin into layers of the object. Consequently, the related AM process, Mask-Image-Projection-based Stereolithography (MIP-SL), can be much faster than the laser- based SLA process by simultaneously forming the shape of a whole layer. An illustration of the MIP-SL process is shown in Figure 1. In the MIP-SL process, a 3D CAD model of an object is first sliced by a set1/17846of horizontal planes. Each thin slice is converted into a two-dimensional (2D) mask image. The mask image is then sent to the DMD and projected onto a photocurable resin surface to form the shape of the related layer.

Each thin slice is converted into a two-dimensional (2D) mask image. The mask image is then sent to the DMD and projected onto a photocurable resin surface to form the shape of the related layer. By repeating the process, 3D objects can be fabricated on a layer-by-layer basis.1.1 Building Speed Limitation of the MIP-SL ProcessIn the MIP-SL process, the building time of each layer consists of spreading liquid resin into a uniform thin layer and curing the formed liquid layer into a solid layer. Compared to a laser beam that is used in the SLA process, the DMD used in the MIP-SLA process can dramatically decrease the curing time of a layer. Hence, the bottleneck for achieving a fast building speed is the spreading of liquid resin into uniform thin layers, which is the focus of the paper.Research systems (e.g. [2-8]) and commercial systems (e.g. [9, 10]) have been developed before based on the mask image projection approach.

[2-8]) and commercial systems (e.g. [9, 10]) have been developed before based on the mask image projection approach. Most of the developed systems are based on the top-down projection as shown in Figure 1. Suppose dLT is the layer thickness. After a previous layer has been cured, the platform in such a system usually moves down a certain distance d and then up by d-dLT in order to spread liquid resin into a uniform thin layer. In addition to the Z movement, a recoating process is usually required to sweep through the platform such that the top surface can be flattened. For resin with low viscosity, a deep-dip recoating approach has also been developed to replace the surface sweeping approach. After the up and down movements in the Z axis, a sufficient waiting time is required for the liquid resin to settle down into a flat surface. However, such recoating methods are time consuming, which limits the building speed of the MIP-SL process.

However, such recoating methods are time consuming, which limits the building speed of the MIP-SL process. Consequently, the building time of such MIP-SL systems is still measured in hours.1.2 ContributionsTo address the building speed limitation of the MIP-SL process, we present recoating approaches for quickly spreading liquid resin into uniform thin layers. For macro-scale parts, our approach is based on a two-way movement design in a bottom-up projection system. We addressed the related challenge of large attaching forces in the bottom-up projection system. By optimizing the process settings, we illustrate that the preparation of a uniform thin layer can be done within seconds.

By optimizing the process settings, we illustrate that the preparation of a uniform thin layer can be done within seconds. Consequently, the developed fast MIP-SL process can build moderate sized parts in minutes instead of hours.Z stagePlatformTankFixtureResinComputerLensMirrorFigure 1: An illustration of the MIP-SL process.Figure 2: An illustration of the bottom-up projection system.2 BOTTOM-UP PROJECTION BASED MIP-SL SYSTEMIn addition to the top-down projection approach as shown in Figure 1, another projection approach used in the MIP-SL process is the bottom-up projection as shown in Figure 2. That is, the mask image is projected onto the bottom of a transparent tank. After a layer is cured at the bottom of the built part, the platform is moved up and then down to form a small gap with the bottom surface of the resin tank.

After a layer is cured at the bottom of the built part, the platform is moved up and then down to form a small gap with the bottom surface of the resin tank. A uniform thin layer can be achieved after the formed gap is filled with liquid resin.2/178472.1 Advantages of the Bottom-up Projection SystemA bottom-up projection based system has several advantages over a top-down projection based system. (1) The container depth is independent of the part height. Thus a shallow vat can be used to reduce the required volume of the liquid resin. During the building process liquid resin can be added by a pump when needed. (2) Recoating is achieved by constraining liquid resin between the previously cured layers and the resin tank. Hence no additional sweeping is needed for flattening the resin surface. (3) Much smaller layer thickness can be achieved since the gap size is only determined by the Z stage resolution regardless of the fluid properties of liquid resin.

(3) Much smaller layer thickness can be achieved since the gap size is only determined by the Z stage resolution regardless of the fluid properties of liquid resin. (4) The curing of liquid resin is sealed from the oxygen-rich environment. By eliminating the oxygen inhibition effect, the liquid photopolymer resin can be cured faster.2.2 Challenges of the Bottom-up Projection SystemDespite the advantages, the bottom-up projection based system has not been widely used in the SLA and MIP-SL processes. A main reason is that the separation of the cured part from the tank surface is difficult. That is, in the bottom-up projection based MIP-SL process, a cured layer is sandwiched between the previous layer and the resin vat.

That is, in the bottom-up projection based MIP-SL process, a cured layer is sandwiched between the previous layer and the resin vat. The solidified material may adhere strongly to the corresponding rigid or semi-rigid transparent solidification substrate, causing the object to break or deform when the build platform moves up from the vat during the building process.One approach to prevent the detachment of a cured layer from the built part is to increase its exposure such that the cured layer can strongly bond to the previous layers. However, such over-curing will also lead to poor surface quality and inaccurate dimensions. Another approach to address the problem is to apply a certain type of coating on the resin vat to reduce the attachment force of a cured layer . Suitable coatings, including Teflon and silicone films, can help the separation of the part from the vat [11, 12]. A coated Teflon glass has also been used in the machines of Denken [13] and EnvisionTec [9].

A coated Teflon glass has also been used in the machines of Denken [13] and EnvisionTec [9]. However, even with the intermediate material, the separation force can still be large. Huang and Jiang [12] investigated the attachment force for the coating of an elastic silicone film. Based on a developed on-line force monitoring system, test results indicate that the pulling force increases linearly with the size of the working area. Experiments indicate that, for a square of 60×60mm, the pulling force to separate the part from the film is greater than 60 N. Such a large attachment force between the cured layer and the vat is a key challenge that needs be addressed in the bottom-up projection based MIP-SL process.To reduce the large attachment force, another approach developed by EnvisionTec in its Perfactory Systems [9] is to incorporate additional mechanisms that add tilting motions in the part separation.

That is, during the separation, one side of the platform can be moved up slowly before the other side. In this way, instead of the pulling-up, the part can be peeled off from the vat surface. Hence the detaching force would be significantly reduced. However, such additional tilting motion will also reduce the building speed of the MIP-SL process.We address the large separation force and the related speed problem for large area parts by developing a two-way movement method. For the purpose of easy sliding between the built part and the resin tank, we use another type of coating material, Polydimethylsiloxane (PDMS, Sylgard 184, Dow Corning). The coating selection is based on the ability of PDMS in inhibiting free-radical polymerization near its surface, as shown by Dendukuri et al [14]. In their research, it was identified that a very thin oxygen-aided inhibition layer (~2.5 μm) is formed that can prevent the cured layer from attaching to the PDMS film.

In their research, it was identified that a very thin oxygen-aided inhibition layer (~2.5 μm) is formed that can prevent the cured layer from attaching to the PDMS film. Thus cured layers can easily slide on the PDMS film. Based on the approach, a fast MIP-SL process has been developed that can build a CAD model in minutes.The remainder of the paper is organized as follows. Part separation forces are analyzed in Section 3. A two-way movement approach to reduce the part separation force is presented in Section 4. Physical experiments on separation forces based on the PDMS film are presented in Section 5. The process settings and the related building time analysis are presented in Section 6. The experimental setup for performing physical experiments is discussed in Section 7. The experimental results of multiple test cases are presented in Section 8.

The experimental setup for performing physical experiments is discussed in Section 7. The experimental results of multiple test cases are presented in Section 8. Finally, conclusions with future work are given in Section 9.3/178483 SEPARATION FORCES OF THE BOTTOM-UP PROJECTION SYSTEMA simple and intuitive approach for separating cured parts from a film is to directly move the platform up a certain distance d and then down by d-dLT where dLT is the layer thickness. Another approach is to move the cured parts horizontally by a certain distance D. The dominant forces that exist in such two approaches are analyzed and compared as follows.3.1 Separation Force in the Up-Down MovementAs shown in Figure 3a, when the platform is moving up, a small gap is formed between the cured part and the film. The resin flowing into the gap is determined by parameters such as the gap h, the surface area, the resin viscosity, and film surface tension.

The resin flowing into the gap is determined by parameters such as the gap h, the surface area, the resin viscosity, and film surface tension. The pulling-up force is a combination of the viscosity and pressure difference that can be expressed as:F z=F F +vPc+FPd=∫v ∂ ( μ x ∂) dz∫ + ΔP dh c∫ + ΔP dx d,(1)where μ is the dynamic viscosity coefficient,cpΔ is the capillary pressure drop across the free surface,anddpΔ is the pressure difference between the inlet and the outlet.

To simplify the problem, we assume that: (i) the problem can be considered as two dimensional and symmetric; (ii) both film and cured surfaces are flat; (iii) the temperature fluctuation during the building process can be omitted; (iv) the tank is big enough such that the flow in the X direction can be considered as infinity; (v) the film can be considered as a fixed bottom wall.

When the areaboundary; while FPd is determined by the size of surface area. In addition,n A n ,PdF∝1(7)A is relatively large, FPd is dominant in Fz. Hence,F z≈FPd∝n A n ,> . (8)1Maximum Separation Force Before Breaking:In the up-down moving process, the main separation force is in the Z direction. As shown in Figure2F=σ t×b=3b, the maximum separation force that can be added to the cured part is:, tσ is the tensile strength of the material and b is the beam cross section size. That is, the maximum_Max zwhereσ t×Aseparation force before breaking follows:F_Max zA∝ . (9)3.2 Separation Force in the Horizontal MovementAs mentioned before, there is a very thin oxygen-aided inhibition layer formed near the PDMS film. When sliding the cured part in the X direction, liquid resin will fill in the gap.

When sliding the cured part in the X direction, liquid resin will fill in the gap. As shown in Figure 4a, the main separation force during the sliding process is the shearing force related to the resin viscosity. That is,F x≈F v⎛ ∫ = ⎜ ⎝v ∂ μ z ∂⎞ ⎟ ⎠dx.(10)Hence,xFA∝ .(11)VxXbA‐ACured partAirbAALLiquidFvFXhzx(b)PDMS Film(a)stationaryFigure 4: An illustration of the separation force during the sliding process.5/17850Maximum Separation Force Before Breaking:As shown in Figure 4b, the maximum bending force allowed for a beam during the sliding process is:F B=b σ × b L 63,wherebσ is the flexural strength of the material, b is the beam cross section size, and L is the length ofthe beam. That is, the maximum separation force before breaking follows:F_Max xA∝3/2.

That is, the maximum separation force before breaking follows:F_Max xA∝3/2. (12)3.3 Separation Force Comparison of the Two MovementsAs shown in Equations (8), (9), (11) and (12), the separation force and the maximum allowed forcesin the Z and X directions have different relationships with the surface area.n A n ,zF F xF Accordingly, we know: (1) For an increasing part size A, the separation force Fz increases much faster than the maximally allowed force FMax_z. In comparison, the maximally allowed force FMax_x increases much faster than the separation force Fx. Consequently, for a large object size, it is desired to use the sliding movement in the X direction.∝> and1FA∝ when the cured part is moved in the Z direction;(a)_Max z3/2A∝ andA∝(b)when the cured part slides in the X direction._Max x(2) For a decreasing part size A, the maximally allowed force FMax_x decreases much faster than the separation force Fx.

In comparison, the separation force FZ decreases much faster than the maximally allowed force FMax_z. Consequently, for a small object size, it is desired to use the up- down movement in the Z direction.Therefore, a new recoating method that is based on the sliding movement in the X direction is needed for the macro-scale MIP-SL process. In contrast, the traditional resin recoating approach based on the up- down movement can be used for the micro-scale MIP-SL process. In the following sections, we will mainly discuss the new recoating method developed for the macro-scale MIP-SL process.4 TWO-WAY MOVEMENT DESIGN FOR THE FAST MACRO-SCALE MIP-SL PROCESSThe experiment results indicate that the suction force between the cured layer and the PDMS film is large during the pulling-up process. Such a large force on the cured layer may cause the building process to fail if the bonding force between the current layer and previous layers is smaller than the suction force.

Such a large force on the cured layer may cause the building process to fail if the bonding force between the current layer and previous layers is smaller than the suction force. In addition, after building multiple layers, such forces on the PDMS film may lead to cracks in the film due to material fatigue caused by the cyclic loading.Based on the PDMS film, we presented a two-channel design for the multi-material MIP-SL process [15]. However, such an approach, mainly designed for switching tanks, is not suitable for the fast building process. In the two-channel design, the building of each layer requires a full cyclic motion including both moving the platform up and down in the Z axis, and moving the tank back and forth in the X axis. Such motions will slow down the building process.

Such motions will slow down the building process. To facilitate a high-speed MIP-SL process based on the bottom-up projection, a novel two-way movement design is developed, which requires much less motions than the two-channel design.An illustration of the fast MIP-SL process based on the two-way movement design is shown in Figure5. In our method, a transparent PDMS film is first applied on the bottom surface of a glass vat.(1) After a mask image is exposed to cure a layer, the platform is moved up in the Z axis for one layer thickness (e.g. 50 μm). Accordingly, the regions of the PDMS film related to the shape of the cured layer will be pulled up by the suction force. However, the force is small due to the super elasticity of the PDMS film. Note that there is no liquid resin between the cured layer and the PDMS film at this moment. (2) The tank is moved along the X axis for a certain distance Δx.

Note that there is no liquid resin between the cured layer and the PDMS film at this moment. (2) The tank is moved along the X axis for a certain distance Δx. A good property of the PDMS film is that a very thin oxygen-aided inhibition layer (~2.5 μm) is formed near the PDMS film that can provide a non-polymerized lubricating layer for easy sliding [14]. If the moving distance is sufficiently large (e.g.6/17851larger than the extent size of the cured layer in the X axis), the elastic deformation of the pulled-up PDMS film will be released by such a sliding movement. Hence, at the end of the X movement, liquid resin will be filled in the small gap between the cured layer and the PDMS film.(3) The mask image of a new layer can now be projected at the bottom surface to cure the next layer. These three steps can then be repeated by moving the tank in an opposite direction. Note that, to achieve the motion in the X direction, we only move the tank and the related frame.

These three steps can then be repeated by moving the tank in an opposite direction. Note that, to achieve the motion in the X direction, we only move the tank and the related frame. There is no relative motion between the platform and the projection device. Hence the XY accuracy of the MIP-SL system will not be affected by the X translations.Step 1: Move up platform △z (1 layer)Z stageCured resinPlatformPDMSLiquid resinLayer thicknessStep 2: Move tank left/right △xZ stagePlatformX stageTankX stageMirrorMirrorLENSLENSStep 3: Projecting mask imageZ stagePlatformX stageTankMirrorLENSFigure 5: The MIP-SL process based on the two-way movement design with PDMS.5 PART SEPARATION EXPERIMENTS BASED ON PDMSPhysical experiments on the separation forces have been presented.

As shown in the test results, the two-way movement design can effectively reduce the large separation force while achieving a fast building speed at the same time.5.1 Separation Force Study based on the Up-down MovementWe studied the part separation force of a cured layer from a coated PDMS glass based on moving the platform up a certain distance d. Since one layer thickness is usually very small (50-200 μm), d is usually much larger in order for the resin to fully fill the gap (e.g. 5mm). In our tests, the PDMS film thickness is set at 1mm. A set of physical experiments have been designed and performed to understand the separation force based on such up-down movement.Figure 6a shows the setup for measuring the pulling-up force. Two FlexiForce sensors (Tekscan, South Boston, MA) with a range of 0-25 lbs are sandwiched between the fixture and the vat. The two sensors are connected to a microcontroller, which can sample and record the sensors’ readouts at over 3KHz.

The two sensors are connected to a microcontroller, which can sample and record the sensors’ readouts at over 3KHz. Since the vat is free at the bottom and the side, and only fixed at the top, the pulling force by the part will be transferred to the sensors when the platform rises. In the experiments, we first use a given7/17852Tankmask image to build a certain number of layers (e.g. 25 layers). The layer thickness is set at 0.2mm. We then begin to record the separation force in the building process of the next few layers.

25 layers). The layer thickness is set at 0.2mm. We then begin to record the separation force in the building process of the next few layers. For each layer, after the designed mask image has been exposed for a certain time, the platform is raised up slowly at 0.6mm/sec for 5mm and the related readouts of the sensors are then recorded.Z stagePlatformFixtureTankPDMSMask image projectionPartResin80 70 60 50 40 30 20 10 0 -10) z o (e c r o F(b)Triangle Square Hexagon Tshape Ushape Band Star(a)Different projection patterns with the same area01040 Position (* 10um)2030506070Figure 6: Separation force study: (a) experimental setup for studying separation forces in the MIP- SL process; and (b) pulling-up forces of a cured layer for T= 0.5 sec, Area = 625 mm2.In our study, we considered three factors that may affect the separation force including (1) exposure time, (2) image area, and (3) image shape.

To understand the effects of these factors, designed experiments were conducted. Seven projection patterns were used for testing the effect of image shape. They are shown in Figure 3, which include circle band, hexagon, t-shape, square, star-shape, triangle, and u-shape. For comparison, all the projection patterns have the same area in the tests. The separation forces of a cured layer were measured based on each of the seven projection patterns. Figure 6b shows the measured separation forces of a sensor for different test cases. The horizontal axis indicates the distance in the Z direction (in the unit of 10 μm), and the vertical axis indicates the measured pulling force (in ounces).

After the cured layer isdetached from the PDMS film, the separation force will drop rapidly from the peak value to 0;(2) Due to the flexibility of the PDMS film, the pulling-up force is rather small within a movingdistance that is less than 200μm.(3) The peak force gets larger when the same mask image is exposed longer; (4) The peak force gets larger when a larger image area is projected; (5) The image shape has more complex effects on the peak force. In addition, their effects mayinteract with the exposure time and the projection area;8/17853(6) With the coated PDMS film on the vat, the separation force is still considerably large (~100 oz or27.8 N for an image area of 625mm2 with 1 second exposure).5.2 Separation Force Study based on the Two-way Movement DesignTo verify the proposed two-way movement design, a set of experiments were conducted based on the setup as shown in Figure 6a. The same set of mask patterns were used in building test layers.

The same set of mask patterns were used in building test layers. The same exposure time and layer thickness were used (1 second and 0.2mm respectively). The building process as shown in Figure 5 was first used in building a set of layers. In the tests, the tank was translated in the X axis by 20mm. The moving speed is set at 25mm/sec. After the layers have been built, the pulling-up forces in the Z axis during building the next layer were recorded. However, instead of curing a new layer as shown in Step 3, the part is moved up slowly at 0.6mm/sec for 2.5mm. The measured forces of a sensor in the Z axis during the aforementioned three steps are shown in Figure 8.

The measured forces of a sensor in the Z axis during the aforementioned three steps are shown in Figure 8. In each figure the curves record the test results based on a sampling resolution of 80 milliseconds.) z o ( e c r o F20 18 16 14 12 10 8 6 4 2 0 -2 -4010Triangle Square Hexagon Tshape Ushape Band Star203040Sampling time (* 20 msec)5060(a)70) z o ( e c r o F20 1816 1412 108 64 202 -4010Triangle Square Hexagon Tshape Ushape Band Star203040Sampling time (x80 msec) Sampling time (* 20 msec)5060Move the platform up by 0.2 mmMove the tank in X by 20 mmMove the platform up by 2.5 mmFigure 8: Pulling-up forces of a cured layer based on the two-way movement design in different settings: (a) T= 1sec, area = 625 mm2; (b) T= 1 sec, area = 156 mm2.The figures show that the force in the Z direction is rather small when the platform is moved up by 0.2mm. During the remaining two steps (i.e.

During the remaining two steps (i.e. sliding on the PDMS film and the platform pulling-up), the peak separation forces are also relatively small (around 2-6 oz or 0.56-1.67 N). Such measured forces are only 3-4% of the related ones as shown in Figure 7. Hence the two-way movement design can effectively reduce the large separation force in the bottom-up projection system.5.3 Shearing Force Study in the X AxisIn the two-way movement design, cured layers can easily slide on the PDMS surface. The FlexiForce sensors were used in a modified setup to measure the shearing force in the X direction. However, no meaningful readouts were recorded from the sensors. To quantitatively estimate the value of the shearing force in the X axis, a set of square rods with different sizes were built using the two-way movement design. The built rods shown in Figure 9 are 10mm tall. The minimum cross section size is 0.4×0.4mm. Note that we also successfully built rods with even smaller sizes.

The built rods shown in Figure 9 are 10mm tall. The minimum cross section size is 0.4×0.4mm. Note that we also successfully built rods with even smaller sizes. However, the rods were so fragile that they lost the mechanical strength to sustain themselves when the part was taken out of the resin vat and washed in isopropyl alcohol.Nevertheless, for a rod with a size of 0.4×0.4mm, the maximum tangential force that can be added on it can be analytically estimated. As shown in Figure 4b, the testing rods in the experiment can be modeled as a cantilever beam. Suppose theFigure 9: Shearing force verification test.9/17854(b)70length of the beam is L, the size of the beam section is b×b, the force in tangent direction is F. The , where I is the section modulus, maximum bending stress at the end can be calculated as:σ =/Mc IIb=4 /12, andcb=/ 2.

The , where I is the section modulus, maximum bending stress at the end can be calculated as:σ =/Mc IIb=4 /12, andcb=/ 2. Substituting these values for their variables, the resultant equation isσ =Suppose the allowable blending stress is []σ and the minimal beam section size is [ ]b . We will have thefollowing equation:F≤3 [ ][ ] b σ L 6. The material used in our tests has the following parameters:]σ =65MPa, [ ]b =0.4mm, L=10mm. According to the equation, the upper bound of the tangential force is [ only 0.07N or 0.25oz. Compared with the separation force in the Z direction, the shearing force in the X direction is rather small.6 A FAST MACRO-SCALE MIP-SL PROCESS AND ITS BUILDING SPEED ANALYSISThe two-way movement design enables the quick spreading of liquid resin into a uniform thin layer. In addition, the DMD-based digital mask projection enables the fast curing of the spread liquid resin into a desired solid layer.

In addition, the DMD-based digital mask projection enables the fast curing of the spread liquid resin into a desired solid layer. Consequently, for a given 3D CAD model, a fast MIP-SL process can fabricate a physical object within a short building time. The curing characteristics and the two-way movement settings of the developed macro-scale MIP-SL process are presented as follows. A detailed analysis of its building time is also discussed.6.1 Curing CharacteristicsThere are two types of photopolymer systems, acrylate chemistry and cationic photopolymerization, in the SLA process [16]. Acrylate chemistry polymerizes via a free-radical mechanism while cationic photopolymerization undergoes ring-opening reactions in the presence of cationic photoinitiators. The monomer propagation for cationic reactions requires relatively higher activation energy. Consequently, the photospeed of acrylate-based photopolymers is higher due to the lower activation energy for free- radical reactions.

Consequently, the photospeed of acrylate-based photopolymers is higher due to the lower activation energy for free- radical reactions. Considering the photospeed difference, we selected the photopolymer resins based on acrylate chemistry for the developed fast MIP-SL process. As shown in Section 6, our projection system can cure a layer within a short exposure time (<500 milliseconds). Such a fast curing time contributes to the desired fast MIP-SL process for building 3D objects in minutes.After an image is exposed for a certain time (Tprojection), a waiting time, Twait_projection, is required before the layer can be moved up for one layer thickness (i.e. Step 1 in Figure 5). Such a waiting time is critical in order for the acrylate resin to complete the solidification process and gain sufficient strength for the Z movement. Otherwise, the building process may fail. The waiting time is dependent on the resin’s curing property.

Otherwise, the building process may fail. The waiting time is dependent on the resin’s curing property. Due to the fast photospeed of the acrylate resins, the waiting time in our system is short (~300 milliseconds in our tests).6.2 Two-way Movement SettingsIn the two-way movement design, the cured part is first moved up for one layer in the Z axis and the tank is then translated in the X axis for a certain distance. The two linear movements have different accuracy and speed requirements.(1) The Z movement needs to be accurate since it will determine the layer thickness of the next layer. The Z stage also needs to have a resolution that is much smaller than a layer thickness. Accordingly, to ensure the desired accuracy and resolution, we set small acceleration and velocity values in the Z movement. The slow movement of the cured part also enables the PDMS film to fully elastically deform for a small attaching force.

The slow movement of the cured part also enables the PDMS film to fully elastically deform for a small attaching force. However, the movement time in the Z axis (TZ) is still reasonably short (~0.4 second in our tests) since only a small moving distance is required (e.g. 50 or 100 μm).(2) The tank needs to be moved in the X axis for a certain distance to release the elastic deformation of the PDMS film. The X moving distance is related to the shape and size of the cured layer, and less than the extent size of the cured layer in the X axis. Since the relative position of the platform and the10/178556FL 3 b.projection system will not change during the X movement, the accuracy and resolution requirements on the X movement are not as high as those on the Z movement.

Hence a much larger acceleration and velocity can be applied in the X movement to reduce the movement time in the X axis (TX).In our testbed, we used a Z linear stage with a thread of 0.5mm/round, and a X linear stage with a thread of 25.4mm/round. The moving time for different displacement distances in our prototyping system was calibrated for both linear stages. After the X movement, another waiting time, Twait_X, is required in order for the flowing liquid resin to settle. Otherwise, the building process may fail. The waiting time caused by the X motions is related to the movement distance and the moving speed. Due to the small gap between the cured part and the PDMS film, the required waiting time is typically short (~100 milliseconds). After the waiting time of Twait_X, the liquid resin forms a uniform thin layer, which is ready for the next layer to be built.

After the waiting time of Twait_X, the liquid resin forms a uniform thin layer, which is ready for the next layer to be built. The process can then be repeated after the related mask image is exposed.Figure 10: The building time of a layer in the two-way movement based MIP-SL process.6.3 The Building Time of a LayerAs shown in Figure 10, the building time of each layer is thus the sum of all the aforementioned steps:TLayer = TProjection + Twait_Projection + TZ + TX + TWait_X.The first two items, TProjection and Twait_Projection, are related to the curing characteristics of the photopolymer resins used in the MIP-SL process. The photopolymer resins based on acrylate chemistry can be quickly cured. A stronger light source used in the projection system can further reduce the projection time TProjection.The other three items, TZ, TX, and TWait_X, are related to the two-way movement design. TZ is related to the layer thickness and the moving velocity in the Z axis.

TZ is related to the layer thickness and the moving velocity in the Z axis. TX is related to the size of the cured layer and the moving velocity in the X axis. A linear stage with a higher speed can be used to further reduce the movement time TX. TWait_X is determined by the gap distance between the PDMS and the cured layer, the moving velocity of the tank, the shape of the cured layer, and the flow properties of the liquid resin. For a typical layer thickness that is usually small, TWait_X is reasonably short (~100 millisecond in our tests).Note that the projection time TProjection for the first few layers is much longer (e.g. 3-4 seconds) to ensure the initial layers can be strongly bonded to the build platform, For all the other layers, the total building time of a layer is usually short (a few seconds in our tests). Hence a fast fabrication speed can be achieved in the developed process (e.g.

Hence a fast fabrication speed can be achieved in the developed process (e.g. building 3mm height per minute).7 EXPERIMENTAL SETUP 7.1 Hardware SystemA prototype system has been built to verify the developed process. The hardware setup of the fast MIP-SL system is shown in Figure 11. In the designed system an off-the-shelf projector (CASIO XJ-S36) was used. The optical lenses of the projector were modified to reduce the projection distance. Various projection settings including focus, key stone rectification, brightness and contrast were adjusted to achieve a sharp projection image on the designed projection plane. The DMD resolution in our system is 1024×768 and the envelope size is set at 48×36 mm. A precise linear stage from Aerotech Inc (Pittsburgh, PA) is used as the elevator for driving the platform in the Z axis. A fast linear stage from Servo Systems Co. (Montville, NJ) is used to drive the tanks back and force in the X axis.

A fast linear stage from Servo Systems Co. (Montville, NJ) is used to drive the tanks back and force in the X axis. A high performance 4-axis motion control board with 28 Bi-directional I/O pins from Dynomotion Inc. (Calabasas, CA) is used for11/17856driving the linear stages. A flat and clear glass Petri dish is used as resin tank. A PDMS film (Sylgard 184, Dow Corning) is coated on the glass dish. 7.2 Software SystemA mask planning testbed has been developed using the C++ language with Microsoft Visual C++ complier. The testbed integrates the geometry slicing and the motion controlling. It also synchronizes the image projection with the X and Z movements.

The testbed integrates the geometry slicing and the motion controlling. It also synchronizes the image projection with the X and Z movements. The graphical user interface (GUI) of the developed software system and the flowchart of the fast MIP-SL process are shown in Figure 12.Z stageGeometry SlicingPlatformControl PanelPower supplySlice geometryResinx stageLoad geometryMove platform to homeMotion ParametersTankBuilding Process ReportENDYesFinished building?NoWait twait_XProjectorLoad next imageControl boardExpose mask imageMove the tank by the current layer’s extent size in the X axisWait twait_projectionMove up platform for one layerFigure 11: The prototype hardware system for the fast MIP-SL process.Figure 12: Flow chart of the fast MIP-SL system and related software system.7.3 MaterialsPerfatory™ SI500 (yellow) and Acryl R5 (red) from EnvisionTec Inc. (Ferndale, MI), were used in testing the developed fast MIP-SL process.

(Ferndale, MI), were used in testing the developed fast MIP-SL process. Both resins belong to Acrylate. For curing depths of 0.05mm and 0.1mm, the exposure times for SI500 based on our projection system are set at 0.3 sec and 0.45 sec, respectively. The exposure times for Acryl R5 are set at 0.4 sec and 0.55 sec for curing depths of 0.05mm and 0.1mm, respectively.8 EXPERIMENTAL RESULTS AND DISCUSSIONTests have been performed to verify the building speed of the developed prototyping system. The results of the designed tests have demonstrated that the presented MIP-SL process can build 3D models in minutes instead of hours.8.1 Test Results of Macro-scale MIP-SLA set of CAD models with different complexity were used in our tests. The screenshots of six input CAD models are shown in Figure 13.a-18.a.

The screenshots of six input CAD models are shown in Figure 13.a-18.a. The related STL files have triangle numbers ranging from several hundreds to 1.2 million (refer to Table 1).Two different layer thicknesses commonly used in the MIP-SL process were tested. A 50μm layer thickness was used in the fabrication of a gear model. The mask image projection time was 0.35 second for each layer except the base. The projection waiting time was set at 0.1 second. For all the other models, a 100μm layer thickness was used in their building processes. Due to the larger layer thickness, a longer12/17857image exposure and projection waiting times were used (0.45 and 0.3 second respectively in the tests). Accordingly the Z movement will also take a longer time for a larger layer thickness.

Accordingly the Z movement will also take a longer time for a larger layer thickness. In the tests, the movement time in the Z axis (TZ) is 0.32 and 0.42 second for the layer thickness of 50μm and 100μm, respectively.The required moving distance in the X axis is related to the size and shape of the cured layer. For a layer with a big cross-section area (e.g. the models of a head and a statue), the X translation distance is set to a value that is close to the X extent size. Due to the large movement, the X waiting time was also set longer. In comparison, for a layer with a small cross-sectional area (e.g. the models of a hearing aid shell and the top portion of a brush), the X translation distance can be much smaller than the extent size of the layer in the X axis. However, due to the fast moving speed in the X axis, the differences on TX are usually small (less than 1 second as shown in Table 1).Two types of resins, SI500 and Acryl R5, were tested.

However, due to the fast moving speed in the X axis, the differences on TX are usually small (less than 1 second as shown in Table 1).Two types of resins, SI500 and Acryl R5, were tested. Their curing characteristics are slightly different. For the same layer thickness, the curing of Acryl R5 takes ~0.1 second longer than that of SI500. The viscosities of the two resins are also slightly different. However, the same settings can be used in the two-way movement design based on the two resins.Figure 12-17 show the built objects based on the developed fast MIP-SL process. The quality of the built objects was examined to be satisfactory. Both surface finish and dimension were analyzed to be acceptable. In our prototyping system, the nominal size of a pixel is 47μm. The fine image resolution enables the meso-scale features (i.e. in the range of 0.1-1 mm) to be well captured in the built physical objects, e.g.

+ Different X moving distances were used in building the bottom (1st -19th layers) and the top (20th -83th layers) portions of the brush.2.52.4head2.32.22.1) c e s (Statue_Rede m T g n dii21.9statuel iu Br e y a L1.81.7Teeth_RedTeeth, Hearing AidBrush_Bot1.6Brush_Top1.5Gear_red1.4Gear1.3051015202530X Movement Distance (mm)Figure 19: Layer building time of the test cases.14/17859All the models shown in Figure 13-18 were built within 12 minutes using our prototyping system. The models with less than 100 layers (e.g. the gear, the teeth, and the brush) only require 2-3 minutes to be built. A statistic of the building time is given in Table 1.A much larger exposure time (e.g. 4-5 seconds) was required for the first few layers in order to build a base. Consequently the built objects and the build platform can be well bonded.

4-5 seconds) was required for the first few layers in order to build a base. Consequently the built objects and the build platform can be well bonded. For all other layers, as shown in Figure 19, the building time of a layer (TLayer) in our MIP-SL process is only 1.4-2.5 seconds. The variation on TLayer is mainly due to different layer thicknesses and the X moving distances. For an average of 2 seconds per layer and a layer thickness of 0.1mm, the building speed of the developed MIP- SL process is ~3mm per minute, or 180mm per hour. To the best of our knowledge, such a MIP-SL process is one of the fastest layer-based additive manufacturing processes that have been developed. A video of building the gear model as shown in Figure 13.a can be found in [17].8.2 Test Results of Micro-scale MIP-SLAs discussed in Section 3, the recoating approach of moving the platform up-down in the Z direction is more appropriate for features that are small.

Accordingly, we have developed a micro-scale MIP-SL prototyping system based on such an approach. The projection area is set at 12.7×8 mm. Various tests have been performed to verify the capability of micro-scale fabrications based on the recoating approach. Two of the test cases are shown in Figures 20 and 21. Figure 20c and 21c show the microscopic images of gear teeth and curved shell with desired thickness (~80μm).

Two of the test cases are shown in Figures 20 and 21. Figure 20c and 21c show the microscopic images of gear teeth and curved shell with desired thickness (~80μm). The support rods as shown in Figure 21b have a diameter around 60μm.(b)(a)1 mm0.2 mm(c)Figure 20: A micro gear: (a) CAD model; (b) Built part; (c) Microscopic picture of the gear teeth.(c)80 um(a)(b)1 mm0.5 mmFigure 21: A micro-scale hearing-aid: (a) CAD model of hearing-aid: (a) CAD model with supports for building job; (b) built part; (c) Micro scope picture of the top view of the built part.15/178609 CONCLUSIONSA novel mask-image-projection-based stereolithography process has been presented for fabricating 3D objects with fast building speed. The proposed approach is based on projecting mask images bottom- up on a PDMS coated glass substrate. A new two-way movement design has been presented for quickly spreading liquid resin into uniform thin layers for macro-scale parts.

A new two-way movement design has been presented for quickly spreading liquid resin into uniform thin layers for macro-scale parts. Such a design can significantly reduce the separation force between cured layers and the resin tank. Experimental results verified that the separation force as well as the sliding force are relatively small during the two-way movement process. The motions related to the two-way movement design can also be performed quickly. The MIP-SL process developed based on such a recoating approach can achieve high fabrication speed for input CAD models.

The MIP-SL process developed based on such a recoating approach can achieve high fabrication speed for input CAD models. The experimental results demonstrate that the newly developed MIP-SL process can successfully fabricate 3D objects with satisfactory quality in a short time (usually in minutes).Some future work to further improve the building speed of the developed MIP-SL process includes: (1) investigating the settings of Tx and Twait_X based on given layers for the best performance; (2) testing faster moving speeds in the X and Z axes; and (3) testing the two-way movement design in prototyping systems with larger XY extent sizes.REFERENCES [1] Bourell, D., M. Leu, and D. Rosen (2009). “NSF workshop - roadmap for Additive Manufacturing:identifying the future of freeform processing.” Washington, D.C., March 30-31, 2009.[2] Chatwin C., M. Farsari, S. Huang, M. Heywood, P. Birch, R. Young, J. Richardson (1998).

from the process.FDM safely generates three dimensional prototypes from 3D CAD software data reducing thenew product development cycle and allowing validation of part design and production tooling.INTRODUCTIONThis paper outlines the use of FDM to speed product design and to stream.line the manufacturingprocess.Buildingprot0l}'pes in tOOlrooDls and model shops is acommonst~t>jnthemanufacturing process, fitting between Part design ··and .actualpartmanufacturing.•••. With.theiadventofCAJj)/CAM systems ••and••the·•. speed••iand••••accuracy·•••these•••••computer..pased.•.•capabilities/bring.·to·.·••thei·desi$Dial1d.·.mallu facturingsteps, it was eXPected that engineers ·would lookiatprototY'ping(as.a logicalnext area in which to •apply advanced technology.The current technological quest has.

The •process builds on early professional experiences with thermofusion control mechanisms and low temperature thermoplastics.FDM is a unique technology to empower design andmMwacturing engineers t<>beable to quickly produce precise, multi-material models in an engineering environment, right at the CAD workstation.The "'#"y:,#,1(!'\1(' ment, measuring 30" xenvironand weighing 250 lbs.Stratasys is a privately held companywith.partial·funding provided by Battery Ventures in Boston. Stratasys MODELERS in the second quarter of 1991.shil)piI1lj,t 3DTHE NEED FOR RAPIDIn today's·business environment, manufacturers need every competitive orh:TOnf'onrAquality product to market as quickly as possible. images created on CAD workstations has become· an additional tool to positively ..............,"""\. and speed (Marks, 1990).ability to rapidly produceBy allowing design and manufacturing engineers to quickly, accurately prototypes, the design process will improve.

and speed (Marks, 1990).ability to rapidly produceBy allowing design and manufacturing engineers to quickly, accurately prototypes, the design process will improve. When an accurate physical model than an hour, the designer can economically create multiple iterations prior toetllcumtJly creategelllerakted in lessRapid prototyping gives shape, form and feel to the image on the COrnPlJlter3D models of complex, sculptured..s111faced parts within minutes or a will increase experimentation and allow improvements to be quickly incorporated (Wohlers, 1990).3D MODELERconjunction with a Stratasys' 3D MODELER is a single step, self-contained modeling system that advantages. Speed is an important benefit of this technology; typical models can minutes rather than hours or days. ThelightweightFDM head operates up to 900 inches (15 inches per second).

Speed is an important benefit of this technology; typical models can minutes rather than hours or days. ThelightweightFDM head operates up to 900 inches (15 inches per second). As no post curing is required, the FDM technique enables the create multiple versions of a part design within a short time frame.The FDM process uses the Stratasys 3D MODELERto116~ Stratasys, Inc.ELEMENTS OF" THE FDM PROCESSIn this process a conceptual geometric model is created on the CAD workstation. into the SOl workstation where it is sliced into horizontal layers that are down-loaded to the 3D MODELER.It is then imported..----1: 0Ir-----lL_·__-,PATTERNLiquid thermoplastic material is extruded and then deposited into ultra thin layers from the lightweight FDM head one layer at a time. This builds the model upward off a fixtureless base.

This builds the model upward off a fixtureless base. The plastic or wax material then solidifies in 1/10 of a second as it is dire¢ted into place with an X-V controlled extrusion head orifice that creates a precision laminate.A spool of .050 inch diameter modeling filament feeds the FDM head and can be changed to a different material in 1 minute.FIXTlIRELESS FOt.lNDAriONI L_~/LAMENr S~~_P_LYMaintaining the liquid modeling material just above the solidification point is fundamental to the FDM process. The thermoplastic melt temperature is controlled to 1 degree Fahrenheit above solidification.

The thermoplastic melt temperature is controlled to 1 degree Fahrenheit above solidification. The 3D MODELER tightly controls the thermoplastic melt temperature just above solidification then employs precision volumetric pumping through the extrusion orifice._30-MOOELERThe Fused Deposition ModeUng (FDM) process produces safe, accurate 3D models in minutes.The model is immediately sheared, quickly solidified, and each layer is bonded to the previouslayer through thermal heating. The completed model is ready to be removed from the 3D MODELER.Successive laminations, within the 0.001 to 0.050 inch thickness range and a wall thickness of.010 to .250 inch range, adhere to one another through thermal fusion to form the model. Our overall tolerance is +or- .005 inches in the X,Y,Z axis over a 12 inch cubed working envelope.117~ Stratasys,SAFETYworkstation.( It stands <5 feettall with a 3' by 3' footprint. The system requires.

used for conceptual mOaellIlginstance, the accuracy of the model allows its use in the. spray Both the investment casting wax and the machinable wax can .,U\"'U-l,lJ.J.....,l'-' by allowing ·the user to go directly to manufacturing using the model· as the malnutacturlng pat:tern.The nylon like filament is a tough m.aterial prodllcingsturdy models suitable for concept models or form, fit and function applica tions.nrr1.I'1f1J~PI'1 by the Stratasys 3DIn the three years the FDM technology was in development several obstacles were overcome. A major breakthrough was the decision to settle on a filament system of material media as opposed toa "hopper" system. The •spool· based filament system has proven· to be a significant strength of the 3D MODELER.The spools give the user the ability to change material in about118i'Stratasys, Inc.one minute by threading the desired material into the prototyping unit.

There is virtually no waste and no vat· to clean out.The materials to produce apart are cost .effective, uSl.lallyunder ten dollars. Materials used withlaser based fabrication systems can cost $20 or more. For example the material for one golfclub head costs approximately $9.00 and one spoolofmaterial can produce roughly 40 club heads.The FDMprocess is not limited by the UV polymers reqUired by many/other rapid prototypingsystems and·newma.terials. are continuously underdevelopment.SUPPORTSThe. FDJv1 l'rocess. does notneed elaborate. supports. to produce a part .as•• dosom.e othersy~t~tns. The 3D. M()DEj14ER .has the ability·· tocreateai.supportiiJrDlid-airratherthanbuiltiing. thesupport.Jlp from••the ·base.· The•••system••is.•also capableofexttudingpl~sticlntofreespacedepending ••ontbepart geometry. When supports are not used,the FDMhead forms a precision horizontal sl.lpport in mid air as it solidifies.OPEN SYSTEMS The 3DM()I.)ELER importsgeollletry thrOl1ghstandard.

The ability to rapidly produce prototypes reduces this source. of manufactur.. ing error or decreases the time it takes to hand produce· prototypes of all components.Th~· aerospace industry.is just one example where fit, form and function area concern.Investment CastingsManufacturing complex metal·componentsrequires a multi..step investment casting process..As. the name implies, the. Stratasys investment casting wax is excellent for lost waxinvestplent castings.minutes.FDMcan stteamlin¢the investment casting process by i allowing the CAI)geometry to investment casting,·using the Sttatasys model as consuming step of hand making the master can be eliminated.In. addition, the shell mold can be rapidly dewaxed using v (K.ennerknecht and Sifford, 1991). The Sttatasys investment casltln~ dally burnt out...This investment casting wax is also machinable. The part can be .U.AAAAOV'....

To be able to hand aclient a prototype of the proposed part at the final presentation or to include a prototype with the proposal package has strong emotional appeal in the sales process. Conceptual modeling also enables engineers to quickly produce multiple iterations of a sample part to streamline the design process.Concept models are important in the shoe industry. One major shoe manufacturer creates hundreds of new shoe heel designs each year. Each heel style is normally produced in one size based on the designer's drawing.

One major shoe manufacturer creates hundreds of new shoe heel designs each year. Each heel style is normally produced in one size based on the designer's drawing. After this initial prototype design is verified, models are created for sizes 3 through 12, with the heel dimensions graded for the size.The ability to quickly and accurately produce the concept model allows this manufacturer toimprove quality and productivity in the manufacturing process.SUMMARYIt is a non-laser based system providing a cost effective, accurate and environmentally safe way to produce 3D models and prototypes.FDM provides a combination of attractive features to provide true desktop 3D modeling.Reducing time to market by accommodating engineering changes quickly and improving productquality demands state-of-the-art prototyping tools.REFERENCESKennerknecht, S.

and Sifford, D., "New Dimensions in Rapid Prototyping Explored for Aluminum Investment Castings," INCAST, March 1990, Vol IV, Number 3, pp. .5 - 10.Marks, P., 1990, "The Rapid Prototyping Revolution...Better Products Sooner," Proceedings, The First International Conference on Desktop Manufacturing, Cambridge, Massachusetts, October, 1990.Wohlers, Terry, "Plastic Models in Minutes," Cadence, July, 1990.For more information contact: Stratasys, Inc.7411 Washington Avenue South Minneapolis, MN .5.5439 Phone: Fax:(612) 941-.5607(612) 944-6741Trademarks:Stratasys, Inc., 3D MODELER, 3D PLOTTER, Fused Deposition Modeling (FDM), Models in Minutes.@ 1991 Stratasys, Inc.122.

ster- Better Molds Through RSP ToolingNew Research and AdvancementsJames R. KnirschRSP Tooling, LLC Reviewed, accepted August 19, 2004AbstractThe recent developments in rapid production tooling have all but made the need for prototype tooling disappear. There are several approaches that are now as fast and inexpensive as prototype tooling, and after part approval can continue to run in high volume production applications. The newest of these approaches is an indirect spray forming process invented by Dr. Kevin McHugh of the Idaho National Engineering and Environmental Laboratories (INEEL). The advantages of RSP Tooling can be found in its accuracy, finish, cost and speed compared to the other rapid tooling processes [1].The commercialization effort for this spray forming process started in February of 2002. The beta production machine was operational in November, 2003, and started to produce production tooling in March, 2004.

The beta production machine was operational in November, 2003, and started to produce production tooling in March, 2004. Since that time tooling has been manufactured and run for many forming applications. In all but the simplest tools the process has proven to be less expensive and faster than standard machining of tools or any other rapid production tooling process. Research and development of the process has continued both at INEEL and at RSP Tooling, LLC making the process faster, more accurate and less expensive to operate. This research has also generated a better understanding of the underlying metallurgy of the process.MethodRSP Tooling is an indirect spray forming technology that was developed by Dr. Kevin McHugh at INEEL for producing molds and dies [2-5]. The general concept (Figure 1) involves converting a mold design described by a CAD file to a tooling master using a suitable rapid prototyping (RP) technology such as stereolithography (SLA).

The general concept (Figure 1) involves converting a mold design described by a CAD file to a tooling master using a suitable rapid prototyping (RP) technology such as stereolithography (SLA). A pattern transfer is made to a castable ceramic, typically alumina or fused silica. This is followed by spray forming a thick deposit of tool steel (or other alloy) on the ceramic pattern to capture the desired shape, surface texture, and detail. The deposit is built up to the desired thickness at a rate of about 500 lb. /hr. Thus, the spray time for a 7” x 7” x 4” thick insert is only 9 minutes. The resultant metal block is cooled to room temperature and separated from the pattern. Typically, the deposit’s exterior walls are machined using a wire EDM, and bolt holes and water lines are added.681Figure 1 - Process SketchThe turnaround time for cavity or insert is unaffected by complexity.

From receipt of a CAD solid model to shipment of the cavity can be a little as 5 days but is typically 10 days. Molds and dies produced in this way have been used for prototype and production runs in plastic injection molding, die casting, and forging operations. Low pressure casting and extrusion dies have been made and are in testing.Generation of the physical model or “master” is straightforward. A number of rapid prototyping (RP) approaches are available commercially to accomplish this, but they differ widely in terms of cost, accuracy, and surface finish.

A number of rapid prototyping (RP) approaches are available commercially to accomplish this, but they differ widely in terms of cost, accuracy, and surface finish. As part of an R&D study conducted with Colorado State University and an industry team [3], the suitability of various RP- generated physical models as well as physical models machined from aluminum and various tooling boards, was assessed for use with RSP Tooling (Figure 2).682Figure 2 - Various Model Methods2More recent research by RSP Tooling, Colorado State University and Cleveland State University have shown that the Stratasys Eden 333 additive method and the Roland MDX- 650 subtractive method can both reduce the cycle time and cost of the initial model by 75% while maintaining or improving the finish and accuracy of the final tool. The nature of the mold design determines which of these methods is the most appropriate.

The nature of the mold design determines which of these methods is the most appropriate. For example a tool with undercuts or deep features would be best run with an additive method, while reasonable flat designs would be easier with the subtractive method.Comparison runs on the same tool were made using the 3D Systems 5000 and the Stratasys Eden 333. Accuracy of the XY dimensions are comparable while the Z axis of the Eden 333, because of its much smaller step size, is significantly more accurate. The flat surfaces of the SLA are better because the Eden 333 shows print lines. The angled surfaces of the Eden 333 are better because of the smaller Z step. Finishing times are approximately the same for both.

The angled surfaces of the Eden 333 are better because of the smaller Z step. Finishing times are approximately the same for both. The Eden 333 is the preferred system because it is significantly faster and less expensive than the 5000, while finishing costs and accuracy are comparable.Ceramic patterns are made by slip casting or freeze casting ceramic slurry, typically made of alumina or fused silica, onto the tool master. This involves mixing a ceramic powder with a liquid activator or binder, pouring the mixture into a mold, allowing it to set up, demolding and firing of the ceramic in a furnace. Many ceramic formulations have been and are being evaluated for suitability in the process. Ease of casting, material cost, surface roughness, strength, thermal shock resistance, maximum use temperature, flatness, and dimensional accuracy are assessed.

Ease of casting, material cost, surface roughness, strength, thermal shock resistance, maximum use temperature, flatness, and dimensional accuracy are assessed. With the right equipment and procedures, very accurate and reproducible ceramics are easily made.Forming the ceramic pattern is one of the most important steps of the process. It represents a significant variable in the cost, timing, finish, and accuracy of the final tool. The ceramic forming process is the most time consuming element in the cycle representing 40% of the entire turn-a-round time for a tool. Research into using a machined graphite or machined ceramic master has been performed and several tests have shown these to be feasible solutions that could eliminate the majority of the time consumed by this step. There are still several hurdles to overcome, and the approach is only feasible for “one of a kind” tools since the master gets destroyed in the process.

There are still several hurdles to overcome, and the approach is only feasible for “one of a kind” tools since the master gets destroyed in the process. No production tools have yet been made using either of these techniques, but development is continuing.The spray forming step is at the heart of the RSP Tooling process. Spray forming involves atomizing: breaking up a molten metal stream into small droplets, using a high velocity gas jet. Aerodynamic forces overcome surface tension forces within the melt producing an array of droplet sizes that are entrained by the jet and deposited onto the pattern (Figure 3).6833Figure 3 - Machine ProcessAs the droplets traverse the distance separating the atomizer and ceramic tool pattern, they cool at very high rates that vary depending on size. As a result, a combination of liquid, solid, and “slushy” droplets impact the ceramic, and “weld” together to form a coherent deposit.

As a result, a combination of liquid, solid, and “slushy” droplets impact the ceramic, and “weld” together to form a coherent deposit. Figure 4 demonstrates the effect of rapid solidification with molten tin sprayed on a party balloon.Figure 4 - Rapid Solidification in ActionThe high cooling rate of the deposit greatly impedes atomic diffusion, so segregation is very limited compared to cast metal. It also minimizes the erosive interaction of the metal and ceramic tool pattern, allowing the deposited metal to accurately capture surface details of the ceramic that would not be possible if the metal were to be cast onto the ceramic.6844The rapid solidification rate also results in non-equilibrium solidification, extendedsolid solubility, and very limited segregation [6] (Figure 5 and 6).Figure 6 - A2 as Deposited at 50xFigure 5 - H13 as Deposited at x500Machine DevelopmentThe Beta machine was completed in January of 2003.

It took 14 months to reach the operational and cost parameters that were established at the beginning of the project. This development work focused on three different areas, machine reliability, ceramic performance, and tool quality. In the beginning the machine was prone to component failure which resulted in molten metal pouring into a sand pit below the nozzle. Most internal components assuming they held up for a run had to be replaced prior to the next run. Costs of the components were high and the cost per shipped tool was astronomical. The challenge was that the internal components run at temperatures of around 1600C. This is also the failure point of many of the materials. We also found that many “experts” were very knowledgeable concerning the standard operations for their products but had little understanding of the underlying science and therefore were of little help with a new application.

Over the next 10 months virtually every internal component was changed to either a new material or better, less expensive supplier. By February of 2004 the machine uptime improved dramatically and component cost dropped to less than $50 per tool.The ceramics used for patterns were also an issue. Ceramics are excellent for high temperature applications but they do not like sudden changes in temperature. In the RSP process the room temperature ceramic is heated to temperatures approaching the melt temperature of steel in less than an hour. Then molten steel is sprayed on to the ceramic. This is brutal treatment for this type of ceramic and most experts indicated it could not be done. A method also had to be developed to mount the ceramic on to the manipulator in such a way it could be easily be loaded and unloaded but still hold up to 100# of steel while being moved in 5 axis of motion.

Initially the ceramics cracked so that they fell off or at best cracked and left deep fissures on the tool face. Through modifications to mixing and firing techniques a6855ceramic that could stand the rapid temperature change was achieved. Mounting systems were also modified so the ceramic could handle the weight and movement.Once the machine was stabilized so that each run was consistent and the ceramic was dependable, work started on the surface and internal quality of the tool. Quality is a thermal issue. It is important to get the metal applied quickly in order to get a porosity free tool but if it is too hot sink holes can result. The first step in development was to achieve a consistent spray. This was achieved through the developments listed above and by redesigning the crucible so that it could be pressurized.

The first step in development was to achieve a consistent spray. This was achieved through the developments listed above and by redesigning the crucible so that it could be pressurized. Once the spray was stabilized it was a matter of trial and error until the right application rate was understood.The machine is now capable of long term continuous operation at a minimal cost per tool and well within the initial cost targets. Research is now directed toward improving the various steps of the operation to improve speed and accuracy still further and to expand the process applications by developing technologies, like bimetallic tooling.AccuracyDimensional accuracy and repeatability of all processing steps have been analyzed by Colorado State University personnel and industry partners using coordinate measuring machines (CMM). This has helped to identify suitable materials and processing conditions.

This has helped to identify suitable materials and processing conditions. Several conclusions have been drawn from the study:Molds made from the same master but different ceramic patterns were essentially identical which is of major importance in multiple cavity dies or replacement inserts. It also means that accuracy can be increased by making a test tool and then modifying the model to the dimensional data.Most of the shrinkage comes from casting and firing of the ceramic. Some ceramic formulations nearly eliminate this shrinkage.

Some ceramic formulations nearly eliminate this shrinkage. It has also been demonstrated that modest variations in binder and firing temperatures have virtually no effect on this shrink.The shrinkage does vary by feature; however this shrinkage is consistent and reproducible which means that a computer program to predict shrinkage by feature will improve the process accuracy.The overall conclusion of the dimensional accuracy study is that the accuracy of molds made by the RSP Tooling method is comparable to the conventional practice of machining, benching, polishing and heat treating. Detailed algorithms are under development that will automatically apply scaling factors to a CAD drawing of an insert for various processing sequences.6866Accuracy has been improved by making a significant reduction in the ceramic shrink rate and by use of the more accurate rapid prototyping model methods.

Further improvements can be expected if machined ceramic or graphite is used.ReplicationThe process can replicate very small features. When sprayed on quartz glass the process reproduced a fingerprint in steel that was accidentally left on the pattern (Figure 7). In tests making a small stamping die with engraved details, features as small as .003 inches could be transferred to the ceramic and then to the sprayed steel tool. All of the detail from a laser burned model was transferred to a tool (Figure 8). This is of even more significance now that the latest SLA machines can achieve details as small as .005 inches in width.Figure 8 - Leather Grain SurfaceFigure 7 - Finger PrintWork with Complete Surface Technologies in Detroit, Michigan has resulted in a process in which grained surface textures can be applied to the freeform fabricated master and then sprayed directly into the tool which eliminates the need to etch the tool after machining.

This again will reduce the time and cost of getting a tool into production (Figure 9).6877Figure 9 - Leather detailTechnical and Economic BenefitsThe main benefits of RSP Tooling involve cost and turnaround time reductions without sacrificing quality or accuracy. When the atomized spray covers the surface of a ceramic tool pattern, it replicates the features very accurately, regardless of the complexity (Figure 10). By so doing, it eliminates many steps in normal mold-making practices such as milling, sink EDM, benching, polishing, and engraving.Figure 10 - Tool DetailSince cost and production time associated with the spray form process are not affected by complexity, the savings achieved by using this process depends on the cost of the alternative process. On recent projects the cost savings varied from $0 to $3000 per insert. Savings in time were proportional.6888The time required to produce multiple cavities is always significantly improved.

On recent projects the cost savings varied from $0 to $3000 per insert. Savings in time were proportional.6888The time required to produce multiple cavities is always significantly improved. The first cavity takes five days but each subsequent cavity follows in 3 hour intervals. This means 32 virtually identical cavities can be shipped in 10 days.The cost for an RSP tool is also a constant, except for the costs of the solid model and the material used. The cost of the master pattern varies depending on the method used, the size, and the material. With the improvements in the model manufacturing methods discussed earlier, both the cost and production time will be dramatically affected. With the right ceramic process, it is conceivable that the delivery time of an insert can be reduced to 8 hours.Die MaterialsThe RSP Tooling machine is designed so each tool can be made from a different alloy.

Because of the rapid solidification of the metal the quality of the tool is the same or in some cases better than machined tools of the same alloy. P-20 when sprayed typically has a better grain and density than the standard machined tool but has the same strength and hardness. Tool life is generally the same.Spray formed H-13 tool steel has significantly better properties than can be achieved by standard methods. As sprayed the tool has a Rockwell C hardness of 56. This can, through artificial aging, be increased to 62Rc. When H13 is aged the strength of the material is significantly increased (Figure 11).Figure 11 - H13 after Artificial AgingThis means that if high wear materials are used in plastic injection molding or if the process is die casting tool life can be extended by 25% or more over a standard H13 tool.

Since the machine can use any metal with a melt point below 1500 degrees centigrade tool life can be further improved by using new or existing hard to machine alloys.6899Dr. McHugh has a DOE grant which is funding research into new alloys which take advantage of the rapid solidification process to increased die life even more. RSP Tooling is investigating new materials such as NiAl which are too hard to machine to have been practical in the past but now can be made into tools at a nominal cost.The density of a sprayed tool is equal to a tool using forged machined steel of the same specification. Tests with H13 have shown densities of 99.5 % to 99.7% [6].

Tests with H13 have shown densities of 99.5 % to 99.7% [6]. This means that not only can water lines be machined with out fear of leaking but the material can be hand polished with out fear of opening up porosity (Figure 12).Figure 12 Polished to Three FinishesA major distinction between this process and other RT processes is that the material used is the same as in a normal machined tool. Thus standard operational practices such as welding, stress relieving, and polishing can all are performed identically (although not as often). To demonstrate that the material is essentially identical tests were performed with sprayed and forged H13 both heat treated identically. Differential thermal analysis (DTA) is useful for analyzing the heat absorption or evolution that accompanies phase transformations, precipitation of new phases, resolution of phases, etc. DTA was performed on samples of spray-formed H13, spray formed/annealed H13 and commercial forged/annealed H13.

DTA was performed on samples of spray-formed H13, spray formed/annealed H13 and commercial forged/annealed H13. Prior to testing, all samples were austenitized at 1015OC and air quenched. Scan data and peak assignments are summarized in Figure 13. Similar peaks corresponding to the exothermic precipitation and growth of carbides, the curie transformation, iron transformation, and melting were observed for all samples.theferrite-to-austenitetransformation,gamma-to-deltaImpact energy analysis was conducted at room temperature on spray-formed H13 tool steel in the as-deposited condition, following artificial aging, and following conventional heat treatment. Various deposition rates and quenching media were used in an effort to modify carbide formation and growth and the nature of the matrix phase, both of which strongly influence toughness.

Various deposition rates and quenching media were used in an effort to modify carbide formation and growth and the nature of the matrix phase, both of which strongly influence toughness. Spray-formed samples were produced at a deposition rate69010of ~250 kg/hr followed by cooling in still nitrogen, at ~250 kg/hr followed by oil quenching, and at ~100 kg/hr with still nitrogen cooling. The impact energy data for spray-formed samples is reflective of bulk rather than surface microstructural features and was strongly influenced by the amount of free carbide and the phases present in the surrounding matrix (martensite and bainite) Microstructural evaluation indicated that spray-formed material was somewhat more refined at the die surface than in the interior [6].

duetothehighercoolingratesFigure 13 - DTA scans on austenitized (1015OC) and air quenched H13 tool steel samples of commercial forged/annealed (Comm-HT), spray formed (SF-HT), and spray-formed/annealed (SF/ANN-HT) materials.Surface FinishThe surface finish that can be achieved is dependent on the ceramic pattern and the initial model. The spray system replicates the ceramic with extreme accuracy and can pick up details as small as .0001 inches. The standard process now in use can achieve a surface finish of 40 micro inches.Several ceramics now being examined have the potential of significantly improving this finish.

The standard process now in use can achieve a surface finish of 40 micro inches.Several ceramics now being examined have the potential of significantly improving this finish. For extremely demanding applications, a fused quartz glass pattern could be sprayed which can result in a mirror finish (Figure 14).69111Figure 14 - Reflection of a PennyCycle Time ImprovementAn additional potential benefit of the spray-forming approach involves the ability to include conformal cooling lines in a tool in order to rapidly cool the molding surface of a mold or die. Mold cooling accounts for about two thirds of the total cycle time in plastic injection molding, die casting and most other metal-mold casting operations. Ideally, cooling lines would be placed near the surface of a die, and would conform to the geometry of the die surface.

Ideally, cooling lines would be placed near the surface of a die, and would conform to the geometry of the die surface. This is referred to as “conformal cooling” and is viewed by molders as very beneficial because it provides better thermal management of the tool and reduced part cycle time. In plastic injection molding for example, conformal cooling has been shown to reduce part cycle time 15-50% compared to standard cooling practices. This improves productivity and reduces energy use [7, 8].The incorporation of cooling lines has traditionally involved machining straight- bore holes into the back of the die insert. Unfortunately, conformal cooling lines can not normally be incorporated into machined dies due to their complex geometries.Dr.

Unfortunately, conformal cooling lines can not normally be incorporated into machined dies due to their complex geometries.Dr. McHugh at INEEL is working under a DoE grant to perfect a dissolvable core that would be inserted into the spray automatically and after the tool was finished the core would be removed resulting in conformal cooling lines (Figure 15).69212Figure15 - Conformal CoolingSeveral experts have suggested that the cycle times of both plastic injection molded parts and metallic die-cast components can be substantially reduced by first depositing a high-strength wear resistant layer (H-13 Steel), followed by a high-thermal conductivity material ( copper) [9,10,11]. Preliminary results indicate that spray forming is well suited for producing these clad tools. A copper/steel clad was formed by depositing a high conductivity copper backing onto an H13 die insert (Figure 16).

A copper/steel clad was formed by depositing a high conductivity copper backing onto an H13 die insert (Figure 16). This allows for a simple water line to be added through machining into the copper cladding yet the cooling will be uniform over the entire surface of the die. Copper backed tools have been manufactured and have shown to have an extremely strong mechanical bond due to the rough surface that is left when only a thin layer of steel is sprayed. Tests are now being run to determine the amount of additional heat that can be removed in this manner and to determine how long the bond holds up to thermal cycling.Figure 16 - Clad Tooling69313LimitationsThere are limitations to the size of molds and dies that can be produced with the current equipment. The original laboratory-scale equipment at INEEL can produce inserts that are about 3” x 3” x 2” thick. Commercial beta equipment located at RSP Tooling, LLC has increased this to 7” x 7” x 4” thick.

The original laboratory-scale equipment at INEEL can produce inserts that are about 3” x 3” x 2” thick. Commercial beta equipment located at RSP Tooling, LLC has increased this to 7” x 7” x 4” thick. However, the process has no inherent size limitation, and machines with larger capacity are being planned. Preliminary research has demonstrated that multiple spray heads can be used within the same machine. This will be the approach used in larger machines. The next proposed size will be 20x20x10 and the machine should be able to produce a tool in less than eight hours.The second limitation is the aspect ratio for standing (boss) features of the mold. Cavity features on the mold surface do not present problems. However, boss features on the mold surface do. Recent projects have shown that this limitation is more significant than originally thought.

However, boss features on the mold surface do. Recent projects have shown that this limitation is more significant than originally thought. For small features the process can now make features with aspect ratios of about 4 to 1 but for larger features (2”x2”) the ratio is closer to 1 to 1. This is because, when spraying molten metal down into a cavity in the ceramic the metal will tend to bridge across the hole before it is entirely filled. While R&D in this area continues, it is currently recommended that these types of features be inserted.ConclusionsSpray forming has demonstrated great potential for reducing the cost and lead time for tooling by eliminating many of the machining, benching and heat treatment unit operations. In addition:Spray forming provides a powerful means to control segregation of alloying elements during solidification, carbide formation and growth, and the ability to create beneficial metastable phases in many tool steels.

As a result, relatively low temperature artificial aging heat treatment can be used to tailor properties such as hardness, toughness, thermal fatigue resistance, and strength, which will increase tool life.Clad tooling with high heat conductivity metals will substantially increase cycle times in plastic injection molding and die cast operations.Successful production runs in plastic injection molding, aluminum die casting, and forging have been demonstrated.Tools are also being made for use in glass forming, extrusion, and permanent molding operations.69414AcknowledgementsThe author gratefully acknowledges support by the Office of Industrial Technologies, Energy Efficiency and Renewable Energy, U.S. Department of Energy under grants DE-FC07-01ID13982 and 1H132 and contributions by Dr. Kevin McHugh, Bruce R. Wickham of INEEL, Prof. James E.

Although there are investigations on the fatigue behavior of SLM-specimens, wide acceptance of SLM is limited by a lack of knowledge concerning the operating behavior of actual parts. This paper presents a review on existing studies about fatigue life analysis in SLM as well as results from uniaxial high cycle fatigue (HCF) tests of 1.4542 stainless steel as-built and machined specimens with a stress ratio of R = 0. Due to a lower surface roughness machined specimens show significantly higher fatigue strength compared to as-built ones. The obtained fatigue strength at 107 cycles of as-built specimens is used as input for fatigue tests of thin-walled actuator housings. Numerical simulation is used to determine the stress distribution of thin-walled as-built actuator housings under specific loads. Results indicate that the thin-walled as-built actuator housing withstand higher peak stresses compared to as-built specimens due to a high stress gradient.1.

Results indicate that the thin-walled as-built actuator housing withstand higher peak stresses compared to as-built specimens due to a high stress gradient.1. IntroductionAdditive manufacturing (AM) is an emerging field in manufacturing technologies that has the common principle of building up solid parts directly from 3D CAD data by adding material layer by layer. Part-specific tools are not required in these technologies. Additive manufacturing technologies allow the generation of parts which fulfil the properties of final products (Gibson et al., 2010; Huang et al., 2013; Zhai et al., 2014). Among AM technologies, selective laser melting (SLM) has become a relevant production method for manufacturing ready-to-use parts made from metals such as stainless steel, nickel, titanium, and aluminum alloys (Gu et al., 2012).

As AM technology SLM enables the following key advantages for innovative products: function optimized and application tailored product design, function integration, mass customization, resource efficiency, and shortened time to market (Grzesiak et al., 2011; Petrovic et al., 2011; Campbell et al., 2012).In SLM process development relevant research has been conducted to increase the parts’ density since it positively affects mechanical properties such as strength and hardness. High density SLM process parameters were developed for various materials such as stainless steel (Kruth et al., 2010), aluminium (Buchbinder et al., 2011), titanium (Kong et al., 2011), and nickel based alloys (Mumtaz et al., 2008). Static material properties of SLM generated parts are comparable to those of conventionally manufactured ones (Zhai et al., 2014).

Static material properties of SLM generated parts are comparable to those of conventionally manufactured ones (Zhai et al., 2014). However, as a powder bed based layerwise process SLM parts are accompanied by a high surface roughness which is a critical factor under cyclical loading (Huang and Leu, 2014). The surface roughness of SLM parts is caused by three main effects: (1) the process inherent stair case effect, (2) partially melted powder particles sticking to the surface, and (3) instability of the melt pool (Thijs et al., 2010; Pyka et al., 2013). Post processing operations such as milling or (electrochemical) polishing allow a smoothening of the surface roughness. Nevertheless, any additional process step is contrary to the direct manufacturing principle underlying AM. In addition, for complex parts with internal features a post processing may not be applicable in any case.

In addition, for complex parts with internal features a post processing may not be applicable in any case. To exploit the SLM technique as single-step and near net shape manufacturing route, a knowledge base about fatigue behavior of SLM generated parts in their as-built condition is required. So far limited studies on fatigueanalysis of as-built specimens or actual parts exist. The lack of knowledge concerning the fatigue behavior of as-built SLM parts under operational loads represents a barrier of entry for industry into SLM technology.The aim of this study is to increase the knowledge base about fatigue life of SLM generated as-built specimens and actual parts. Experiments using as-built and machined SLM 1.4542 stainless steel specimens are used to determine the Wöhler-curves up to 107 cycles. The fatigue strength in the high cycle fatigue (HCF) regime between these two conditions is compared.

The fatigue strength in the high cycle fatigue (HCF) regime between these two conditions is compared. Based on fatigue testing of as-built specimens the fatigue behavior of SLM generated thin-walled actuator housings1 in their as-built condition will be analyzed. To estimate the stress distribution within the thin-walled SLM housings under defined loads a numerical simulation model will be used. Prior to the experiments a comprehensive overview about fatigue analysis in selective laser melting of metal parts is given.2. Review of fatigue analysis in selective laser meltingOver the last decade studies in the field of selective laser melting were primarily focused on process qualification, development of new materials and applications. Examples are studies which investigate the application of SLM for manufacturing medical implants with tailored properties (e.g. porosity) (Murr et al., 2009; Bartolo et al., 2012).

Examples are studies which investigate the application of SLM for manufacturing medical implants with tailored properties (e.g. porosity) (Murr et al., 2009; Bartolo et al., 2012). Only in recent years studies on fatigue analysis of SLM manufactured specimens have been published and set up a further research area.Fig. 1 summarizes the results form fatigue analysis using Wöhler tests which have been conducted by various research groups. Each data point represents the characteristic knee point of the obtained S-N curves towards the long life fatigue regime. The analysis of existing studies (Table 1) allows the following conclusions:The present studies are limited towards the material selection. Most fatigue analysis is focused on stainless steel followed by studies on the SLM well approved Ti6Al4V alloy. Both for AlSi10Mg and Hastelloy® X one study is known.

Most fatigue analysis is focused on stainless steel followed by studies on the SLM well approved Ti6Al4V alloy. Both for AlSi10Mg and Hastelloy® X one study is known. Independently conducted studies on machined 1.4542 (17-4 PH) stainless steel performed by Sehrt and Witt (2010) as well as by Starr et al. (2011) show good accordance. Similarly, results of Rafi et al. (2013) and Spierings et al.

1. Summary of knee points (S-N values) from relevant studies on the fatigue life of SLM manufactured specimens.1 Function integration is a key advantage of AM technologies. A novel process has been developed to structurally integrate piezoceramic multilayer actuators using the SLM technology (Stoffregen et al., 2013). In this process the actuators are fully integrated into a metallic, monolithic SLM housing. Since the actuator induces cyclic loads knowledge about the fatigue strength of the surrounding housing is required.(15-5 PH) are comparable towards the fatigue strength at 107 cycles and the knee point.Studies are primarily focused on the fatigue analysis of post processed SLM specimens. In most cases for post processing a surface finishing such as turning and polishing was applied. In addition, some studies investigate the influence of subsequent heat treatments. Only three studies on as-built specimens, specifically without surface finishing, could be identified.

In addition, some studies investigate the influence of subsequent heat treatments. Only three studies on as-built specimens, specifically without surface finishing, could be identified. Spierings et al. (2013) investigated different surface treatments on fatigue life of 1.4404 (316L) stainless steel. Between polished and machined specimens no significant difference has been found. In contrast, the as-built specimens showed a lower fatigue strength due to the high surface roughness (Ra ≈ 10 µm). Results of Wycisk et al. (2013) on Ti6Al4V specimens showed a significant difference in fatigue life between as-built and polished specimens as well. The fatigue strength of sand-blasted (Gong et al., 2012) and turned (Rafi et al., 2013) Ti6Al4V specimens show similar values to the results of Wycisk et al. (2013) (finished). The number of cycles at the knee points however shows a variance. One reason might be the low sample number of the studies which leads to statistical uncertainty.

(2013) (finished). The number of cycles at the knee points however shows a variance. One reason might be the low sample number of the studies which leads to statistical uncertainty. The research group of Leuders et al. (2013) investigated different heat treatments on the fatigue life of Ti6Al4V specimens. A significant positive influence of subsequent heat treatment was observed. In addition, the application of hot isostatic pressing (HIP) could further extend the fatigue life due to reduction of porosity. A similar research design was applied by Riemer et al. (2014) on 1.4404 (316L) stainless steel specimens. Besides a heat treatment and a HIP process of machined specimens a set of specimens in their as-built condition was tested. Due to the high surface roughness as-built specimens showed the lowest fatigue strength compared to machined ones. The heat treatment had a positive influence on fatigue life; specimens undergoing a HIP process showed best results.

The heat treatment had a positive influence on fatigue life; specimens undergoing a HIP process showed best results. The positive influence of a HIP treatment of SLM generated specimens on fatigue life could be found as well for Hastelloy® X by Wang (2012). For AlSi10Mg the fatigue strength was determined with and without heat treatment (peak-hardening) on machined specimens (Brandl et al., 2012). A significant positive influence of peak-hardening was observed.The reviewed studies predominantly investigated SLM specimens manufactured vertically under a polar angle of 0°. Brandl et al. (2012) studied three different polar angles (0°, 45°, 90°) combined with a variation of heat treatment and platform temperature using a DoE approach. The orientation of the specimens had a significant influence on fatigue strength depending on the other two parameters.

The orientation of the specimens had a significant influence on fatigue strength depending on the other two parameters. Above maximum stresses of 600 MPa Hastelloy® X specimens showed a higher fatigue resistance when horizontally oriented compared to vertically oriented. Below this stress level the difference was not significant (Wang, 2012). A study on Ti6Al4V specimens manufactured in x, y, and z orientation showed qualitatively better fatigue performance for horizontally than for vertically oriented specimens. However, due to porosity and residual stresses this effect was not quantified (Edwards and Ramulu, 2014).The fatigue life of finished SLM specimens is comparable to conventionally manufactured specimens. For 1.4404 (316L) little difference in fatigue strength between SLM and conventional specimens was observed (Spierings et al., 2013). A similar conclusion is provided by Sehrt and Witt (2010) for 1.4540 (15-5 PH) stainless steel. The study of Rafi et al.

A similar conclusion is provided by Sehrt and Witt (2010) for 1.4540 (15-5 PH) stainless steel. The study of Rafi et al. (2013) on Ti6Al4V showed a higher fatigue strength for SLM specimens compared to cast and annealed specimens due to a fine lamellar microstructure resulting from the SLM process. In general, when comparing fatigue strength of SLM specimens with conventionally manufactured specimens the surface and heat treatment need to be considered (Rafi et al., 2013).Studies on the fatigue behavior of SLM specimens were primarily carried out in the high cycle fatigue regime (5 ∙ 104 < N ≤ 107). Table 1 summarizes the identified studies with corresponding cycle limits and stress ratios. Only the study of Brandl et al. (2012) investigated a cycle limit beyond 107.The sample size for fatigue tests is between 5 to 15 specimens in one condition.

Only the study of Brandl et al. (2012) investigated a cycle limit beyond 107.The sample size for fatigue tests is between 5 to 15 specimens in one condition. One exception is the study of Sehrt and Witt (2010) which used 40 specimens made from 1.4542 (17-4 PH) stainless steel.Aside of the tests on fatigue strength, studies on the fatigue crack growth were performed by Riemer et al. (2014) on 1.4404 (316L) specimens and the research groups of Leuders et al. (2013) and van Hooreweder etal. (2012) on Ti6Al4V specimens. In addition, fatigue analysis on additively manufactured polymers were conducted by van Hooreweder et al. (2010), van Hooreweder et al. (2013), Blattmeier et al. (2012), and Lee and Huang (2013).A recent study of Lipinski et al. (2013) is the only known research work on fatigue analysis of actual parts manufactured using SLM.

(2013), Blattmeier et al. (2012), and Lee and Huang (2013).A recent study of Lipinski et al. (2013) is the only known research work on fatigue analysis of actual parts manufactured using SLM. In their work two different sample sets made from commercially pure grade 2 titanium (CPG2Ti) were tested: 1) thin-walled tube specimens with (w) and without (wo) heat treatment; 2) porous specimens (complex 3D lattice structures with geometrical defined porosity characteristic as bone substitutes (Barbas et al., 2012)). S-N curves for thin-walled tube samples with (n = 18) and without (n = 40) heat treatment were obtained using a stress ratio of R = 0.1. The heat treatment showed a negative influence on fatigue limit due to a relaxation of compressive residual stresses. In general, fatigue stress amplitudes between 65.7 MPa (w) and 74.5 MPa (wo) at N = 107 found to be relatively low compared to the CPG2Ti bulk material properties.

In general, fatigue stress amplitudes between 65.7 MPa (w) and 74.5 MPa (wo) at N = 107 found to be relatively low compared to the CPG2Ti bulk material properties. The high surface roughness and SLM specific defects/imperfections were claimed as reasons therefore. For tests on two different porous specimens different cyclic load levels were applied which lead to variations in stress ratio and amplitude. Using this experimental setting a method based on the stress gradient of the structure was developed to define a fatigue criterion for geometrical defined porous structures.Table 1.

So far research in the following areas has been of subordinate importance:Investigation of the behavior of SLM parts in the very high cycle fatigue (VHCF) regime: Load bearing parts in the automotive and aerospace industry are often exposed to cyclic loads beyond 107 cycles. However, the determination of S-N curves in the HCF regime, which is predominantly existent with respect to SLM, does not necessarily allow an estimation of the very high cycle fatigue behavior since the asymptote must not be horizontally (Bathias, 1999; Marines I. et al., 2003; Sonsino, 2007; Pyttel et al., 2011; Stanzl-Tschegg, 2014). The references indicate that the course of the S-N curves is material dependent. When a so-called “endurance limit” is used for design, the safety against a declining S-N curve in the VHCF regime will decrease (Sonsino, 2007).

When a so-called “endurance limit” is used for design, the safety against a declining S-N curve in the VHCF regime will decrease (Sonsino, 2007). Based on applications of SLM manufactured parts with loads beyond 107 cycles relevant studies have to be performed which allow a statistical significant assessment of fatigue strength in this regime.Studies on repeatability of fatigue characteristics: As shown in Table 1 most Wöhler tests were performed based on 6 to 15 samples which is at the lower end of the recommendation of 12 to 24 samples for design allowable and reliability tests (Lee et al., 2005). With increasing sample size the confidence of the fatigue strength can be increased (Beretta et al., 1995; Martin et al., 2011). In addition, the influence of build direction and position as well as inter and intra batch comparisons have not been studied systematically so far.

In addition, the influence of build direction and position as well as inter and intra batch comparisons have not been studied systematically so far. This might be of further interest, since it is known from other studies that the SLM process is accompanied by anisotropy (Tolosa et al., 2010; Chlebus et al., 2011; Spierings et al., 2013; Thijs et al., 2013) which may also affect the fatigue behavior.Studies on actual parts manufactured by SLM under operational loads: Present studies follow a standard protocol using uniaxial or bending tests based on specimens. Test series to assess the fatigue behavior of SLM generated parts also under complex (multiaxial, stochastic and/or operational) stress-strain scenarios have not been conducted. In addition, limited knowledge on the fatigue behavior of as-built specimens (e.g. high surface roughness) is present which is of special interest to use SLM as direct single-step manufacturing technology.3.

high surface roughness) is present which is of special interest to use SLM as direct single-step manufacturing technology.3. Research methodology3.1 System and materialFor the purpose of this study all specimens were manufactured on an EOS M270 SLM system with a 200 W Yb-fibre laser (continuous wave, wave length λ = 1064 nm, Gaussian beam profile), using commercially available 1.4542 (17-4 PH) stainless steel powder (d50 ≈ 28 µm; d90 ≈ 41 µm). The system integrated beam expander was set to a laser spot diameter ds ≈ 140 µm. PRIMES Focus/Beam Monitor is used to verify the laser power output and laser spot diameter. For all specimens, the layer thickness was fixed to 20 µm, representing the default value for the combination of material used and the SLM system. All specimens were manufactured using a meander-shaped stripe exposure which rotates along the z-axis.

All specimens were manufactured using a meander-shaped stripe exposure which rotates along the z-axis. A qualified exposure parameters which lead to fully dense parts (laser power P = 195 W; scan speed vs = 1000 mm/s; hatch distance h = 0.1 mm) was applied.3.2 Uniaxial fatigue test of SLM manufactured specimensAs test geometry for fatigue analysis common round un-notched specimens (Kt = 1.0) according to ASTM E466-07 with tangentially blending fillets between the test section and the ends have been chosen. The nominal diameter was set to 6 mm, the length of the constant test section to 11 mm. All specimens were manufactured vertically oriented. 39 specimens were manufactured in net-shape to be tested as-built, 20 specimens were manufactured with 1 mm allowance and machined afterwards. The as-built specimens were manufactures in two batches. Fatigue tests were performed on an Instron servohydraulic Hydropuls MHF test system at ambient conditions.

The as-built specimens were manufactures in two batches. Fatigue tests were performed on an Instron servohydraulic Hydropuls MHF test system at ambient conditions. A uniaxial test setup with a stress-ratio of R = 0 (pulsating tension) and N = 107 cycles as the maximum number of cycles was selected. The test frequency was varied between 70 Hz and 140 Hz depending on the applied load. In this regime a significant influence of the test frequency is not expected due to the absence of corrosion induced degradation effects and temperature rise of the specimens (Bargel and Schulze, 2005; Spickenreuther, 2007).The routine test (Weibull, 1961) in combination with the staircase method (Lin et al., 2001; Morrissey and Nicholas, 2006) was used to determine the S-N curves and the fatigue strength at N = 107 cycles. The initial stress of 500 MPa for as-built and 600 MPa for machined specimens has been reduced stepwise until no failure occurred.

The initial stress of 500 MPa for as-built and 600 MPa for machined specimens has been reduced stepwise until no failure occurred. Afterwards the staircase tests were performed for as-built and machined specimens. Since the as-built specimens were manufactured in two batches the staircase method has been applied for each batch separately. To determine the statistical distribution of the S-N curve each stress horizon has been tested twice for machined and five times for as-built specimens (three from batch one, two from batch two). Analysis of fatigue tests was carried out with a survival probability of 10%, 50%, and 90% using a Gaussian probability grid and the formula of Rossow (1964). The staircase tests were evaluated based on the method proposed by Hück (1983) which takes account of the step interval and the stress horizon of all valid data points.

The staircase tests were evaluated based on the method proposed by Hück (1983) which takes account of the step interval and the stress horizon of all valid data points. The width of the scatter band has been assumed to be constant after the knee point (Sonsino, 2008).The roughness of three machined and three as-built specimens has been determined using a contactsurface measurement system (Perthometer).3.3 Uniaxial fatigue test of SLM manufactured thin-walled as-built actuator housingsIn previous research a process has been developed to fully integrate piezoceramic actuators into a metallic, monolithic housing by means of SLM. The main advantage is an application tailored housing design, particularly the stiffness of the housing (Stoffregen et al., 2013). Therefore, the housing shown Fig. 5a has three design parameters, wall thickness t, housing radius r, and angle of bellow α, which can be adjusted to obtain proper stiffness characteristics.

Therefore, the housing shown Fig. 5a has three design parameters, wall thickness t, housing radius r, and angle of bellow α, which can be adjusted to obtain proper stiffness characteristics. A numerical simulation model has been used to determine the stress distribution within the housings under specific loads. Three different housings have been tested. The selection was based on the fatigue strength obtained from the as-built specimens and the resulting stress distribution under application relevant loads. Application relevant loads were derived from typical blocking forces of integrated actuators (maximum value 1.5 to 2 kN). Fatigue analysis of the thin-walled housings is referred to the peak stress.The selected housings were tested similarly to the specimens (section 3.2). A uniaxial test setup with a stress-ratio of R = 0 was chosen and corresponds to the load scenario induced by integrated actuators. The maximum number of cycles to be performed was set to N = 107 cycles.

The maximum number of cycles to be performed was set to N = 107 cycles. The test frequency was set to 133 Hz for all thin-walled housings due the low applied forces. For all three housing geometries a routine test method was used. For two housing types 10 samples were tested, for the other one 29 samples. In addition to the routine test the staircase method was used for the latter. Analysis corresponds to the procedure described under section 3.2.4. Results4.1 Fatigue strength of as-built and machined specimensThe arithmetical mean deviation of the roughness profile for as-built specimens is on average Ra = 13.7 µm, the maximum height of the roughness profile is on average Rz = 86.0 µm. For machined specimens the roughness parameters are Ra = 0.2 µm and Rz = 1.7 µm. Fig. 2 shows the results of the fatigue tests of as- built and machined specimens. Corresponding survival probability levels are included for 10%, 50%, and 90% probability.

Fig. 2 shows the results of the fatigue tests of as- built and machined specimens. Corresponding survival probability levels are included for 10%, 50%, and 90% probability. The calculated fatigue strength at 107 cycles is σmax = 219 MPa for as-built specimens, which corresponds to a stress amplitude of σa = 110 MPa. For machined specimens the maximum stress at 107 cycles is σmax = 492 MPa (σa = 246 MPa). Table 2 summarizes the characteristic parameters of the obtained S-N curves. The slope in the finite life region is significant flatter for machined compared to as-built specimens.Fig. 3 illustrates the results of the staircase testing method.

The slope in the finite life region is significant flatter for machined compared to as-built specimens.Fig. 3 illustrates the results of the staircase testing method. A fictive test result has been added for as-105106107100200300400500600700machined P90Maximum stress (max) (MPa)Number of cycles to failure (Nf)Stress ratio R = 0as-built #1 failuremachined run outmachined P10machined failureas-built #2 run outas-built #2 failureas-built P10as-built P90machined P50as-built #1 run outas-built P50Fig. 2. S-N curves with survival probability for as-built and machined specimens (log-log scale).as-built #2 run outas-built #1 failuremachined staircasemachined fictiveMaximum stress (max) (MPa)Specimen numberas-built #2 fictivemachined failuremachined run outas-built #1 run outas-built #1 staircaseas-built #2 staircaseStress ratio R = 01234567891011121314200300400500as-built #2 failureFig. 3.

3. Staircase testing results for as-built specimens (batch 1 & 2) and machined specimens.built specimens (batch 2) and machined specimens according to Hück (1983). For as-built specimens the two conducted staircase tests show a divergence of the stress amplitude in the range of 10 MPa. Since manufacturing parameters have not been changed and significant differences in microstructure and hardness among the specimens from batch 1 and 2 have not been observed, both batches are analyzed together. In addition, previous conducted inter and intra batch comparisons on tensile strength of SLM manufactured 1.4542 stainless steel specimens demonstrated reproducibility.Table 2.

4c). Contrary, for machined specimens a semi-circular shape is observed (Fig. 4a). Prepared micrographs of machined and as-built specimens showed no significant influence of the post-processing on the microstructure (e.g. surface hardening), which may influence the fatigue strength.4.2 Fatigue strength of thin-walled as-built actuator housingsThe fatigue strength results at 107 cycles of as-built specimens are used as input for the fatigue analysis of thin-walled as-built actuator housings. An ANSYS mechanical simulation model of the housings is used to determine the geometrical dependent stress distribution under application relevant loads. The design parameters of the housing are wall thickness t, housing radius r, and angle of bellow α (Fig. 5a). Three different housing designs are tested. The selection criterion is based on operational loads which result in a maximum stress (von Mises) close to the fatigue strength at N = 107 cycles of as-built samples (σmax = 219 MPa).

The selection criterion is based on operational loads which result in a maximum stress (von Mises) close to the fatigue strength at N = 107 cycles of as-built samples (σmax = 219 MPa). Micrographs of the bellow fold are used to determine the as-produced wall thickness (Fig. 5b). The numerically obtained stress distribution displaying position and value of the maximum stress is shown in Fig. 5c exemplarily for housing 2. Since it is assumed that the housing design results in a notch effect which reduces the fatigue strength one initial housing test is performed with an operational load resulting in a lower maximum stress. Due to a divergence between simulated (kh,sim) and experimentally determined (kh,exp) housing stiffness a correction of the test load Ftest following the equation( ) ()is conducted. Using this equation a maximum test load can be calculated from an expected maximum peak stress σmax,p.

Using this equation a maximum test load can be calculated from an expected maximum peak stress σmax,p. An approximate linear relationship between load and maximum stress is shown during FE analysis. Table 3 summarizes the initial test load for the thin-walled housing as well as relevant design1234(a)(b)(c)Fig. 5. (a) Cross section of the thin-walled actuator housing and design parameters. Numbers indicate position of the bellow. Blue marked are inner folds, red marked is the outer fold at which the numerical simulation shows the highest peak stress. (b) Micrograph of a bellow fold (housing 2); black: cross section of the fold, grey: embedding resin. (c) FEA result of stress distribution in an outer bellow fold for housing 2.(1)parameters.Table 3. Initial test setup for thin-walled actuator housings and corresponding housing design parameters.HousingTest load FtestStress (v.

Initial test setup for thin-walled actuator housings and corresponding housing design parameters.HousingTest load FtestStress (v. Mises)*Stiffness kh,simStiffness kh,expDesign parameter (t, r, α)11.85 kN213 MPa93 N/µm84 N/µm0.6 mm, 8.5 mm, 160°20.76 kN213 MPa34 N/µm25 N/µm0.5 mm, 10.0 mm, 137.5°30.82 kN182 MPa44 N/µm35 N/µm0.6 mm, 8.5 mm, 137.5°maximum peak stress ( ) based on FEAFig. 6 shows the results of the fatigue tests of the three different housings designs. The maximum stress (peak stress at notch root) is calculated from the test load using equation (1). The knee point for all tested housings is in the range of 7.5 to 10 ∙ 105 cycles and therewith in a higher cycle regime than the tested as-built specimens. In contrast, the scattering range is below the tested as-built specimens, both horizontally (TN) and vertically (Tσ). The slopes in the finite life region show little variance among the three housings.

The slopes in the finite life region show little variance among the three housings. Compared to the fatigue strength at N = 107 cycles of the as-built specimens the housings withstand a higher maximum stress. When exhibiting failure the cracks occur at one of the three inner folds (Fig. 5a). A predominantly position of the crack among the three inner folds is not observed. Table 4 summarizes the parameters of the S-N curves for the thin-walled housings.5. DiscussionA strong divergence between the fatigue behavior of as-built and machined SLM 1.4542 stainless steel specimens in the HCF regime is observed. For the machined specimens the fatigue strength at N = 107 cycles is 492 MPa and accounts 51% of the measured ultimate tensile strength (Rm = 961 MPa). This finding is in accordance to the studies of Sehrt and Witt (2010) and Starr et al. (2011) which obtain values of 52% respectively 56%.

Mises) σmax,p / Peak stress amplitude (v. Mises) σa,p343 MPa / 172 MPa425 MPa / 213 MPa355 MPa / 178 MPaNumber of specimens291010Number of run outs521Horizontal scattering range TN (arithmetic average)1.311.631.28⁄Vertical scattering range Tσ at N = 1071.071.191.05⁄behavior between machined SLM generated specimens and conventional manufactured ones.For as-built specimens the fatigue strength at N = 107 cycles is with 219 MPa significantly below the value of machined specimens. The fatigue strength accounts only 23% of the ultimate tensile strength. The reason can be found in the high surface roughness of as-built specimens and the therewith associated notch effect (Arola and Williams, 2002; Radaj and Vormwald, 2007; Chan et al., 2013). The surface roughness due to partially melted powder particles and the inconsistency of the melt pool causes micro notches at the specimen’s surface which lead to a stress intensity during cyclical loadings.

The analysis of the fracture surface shows that crack initiation occurs at defects on the specimen’s surface and supports this assumption. Based on the fatigue results an experimentally determined surface roughness reduction factor γ in the HCF regime can be calculated:For as-built specimens a significant divergence between the two batches is not assumed. The following reasoning underlies this procedure: (1) Divergence between the two batches is only present above 9 ∙ 105 cycles. In the finite life region the results indicate random scattering. (2) Manufacturing parameters have not been changed and significant differences in microstructure and hardness among the specimens from batch 1 and 2 have not been observed.

(2) Manufacturing parameters have not been changed and significant differences in microstructure and hardness among the specimens from batch 1 and 2 have not been observed. (3) The vertical scattering Tσ = 1.38 in the HCF regime of the combined analysis of batch 1 & 2 is in the range of cast and sintered steel (Sonsino, 2007).The slope in the finite life region k as well as the cycle number at the knee point Nk of machined specimens is typical for un-notched cast and sintered steel with medium strength (Sonsino, 2007). Machined specimens (slope k = 20) show a higher sensitivity towards stresses in the finite life region compared to as-built specimens (slope k = 3). With increasing notch effect (e.g. surface roughness) the slope becomes steeper and the sensitivity towards stress variations decreases (Lee et al., 2005; Radaj and Vormwald, 2007).

With increasing notch effect (e.g. surface roughness) the slope becomes steeper and the sensitivity towards stress variations decreases (Lee et al., 2005; Radaj and Vormwald, 2007). This finding is true for the as-built specimens which show a significant steeper slope compared to machined ones; in case of as- built specimens crack propagation mechanisms are dominant.The fatigue analysis of the thin-walled housings shows that the maximum stress at the notch root is significantly above the value of as-built specimens (σmax = 219 MPa). The fatigue strength of the housings at N = 107 cycles is between 343 MPa for housing 1 and 425 MPa for housing 2. A potential explanation can be found in the stress gradient. For the parameters shown in Table 3 an FEA is applied to determine the stress(2)distribution at the folds.

For the tested housing the order of the stress gradients is χ1 < χ3 < χ2. This result implies that housing 1 should have the lowest fatigue strength and housing 2 respectively the highest fatigue strength which is in accordance to the obtained S-N curves in Fig. 6. Therefore, besides peak stresses the stress gradient needs to be considered for fatigue analysis of the thin-walled as-built actuator housings.Of practical relevance is the little scattering of the test results for the thin-walled actuator housings. Both, horizontal and vertical scattering, are significantly below the ranges of the as-built and machined specimens. This allows a fatigue life approach with a low safety factor.Based on the FEA results indicating a maximum peak stress at the top outer fold (Fig. 5a, position 4) a failure is assumed to occur at the same position. Contrary when exhibiting failure it occurred at one of the three inner folds (Fig. 5a, positions 1-3).

5a, position 4) a failure is assumed to occur at the same position. Contrary when exhibiting failure it occurred at one of the three inner folds (Fig. 5a, positions 1-3). Crack initiation starts from the outer side from one of the three inner folds. A potential explanation is the geometry of the thin-walled housing. Outer folds are characterized by a larger diameter compared to inner folds while having the same wall thickness. Consequently, outer folds have a larger cross section compared to inner folds which results in a lower average stress. The latter might be the reason why the thin-walled housings exhibit failure at the inner folds.The thin-walled housings demonstrate fatigue strength at N = 107 cycles for loads up to 1.5 kN which fulfils the requirements of the proposed application (integration of piezoceramic actuators). For housing 1 fatigue strength at 107 cycles exists even for loads up to 2.5 kN.6.

For housing 1 fatigue strength at 107 cycles exists even for loads up to 2.5 kN.6. Summary and outlookThe present work shows a review of fatigue analysis in selective laser melting. Within this comparably new research field studies are conducted almost exclusively on the basis of specimens. The studies focus on SLM approved stainless steel (e.g. 1.4542, 316L) or Ti6Al4V alloy which are predominantly tested in post processed conditions (e.g. machining, heat treatment) in the HCF regime (N ≤ 107 cycles); the fatigue behavior of as-built specimens has not been investigated intensively. Therefore, valid material values for the fatigue life approach of as-built specimens and SLM parts do not exist. However, for machined specimens fatigue strength is in the range of conventionally manufactured specimens. Standardized fatigue test conditions and0.00.10.20.30.40.50.60.76080100120140160180200220Stress (v.

Standardized fatigue test conditions and0.00.10.20.30.40.50.60.76080100120140160180200220Stress (v. Mises) (MPa)Distance (mm)Housing 1χ = 3.7χ = 4.9χ = 5.9Housing 2Distance (mm)Housing 3Distance (mm)0.00.10.20.30.40.50.60.760801001201401601802002200.00.10.20.30.40.50.60.76080100120140160180200220Fig. 7. Stress distribution (v. Mises) of the three different housings under the loads shown in Table 3. Stress distribution is shown in radial direction using FEA. Stress gradient is calculated for notch root (x = 0 mm). Dashed line represents polynomial fit of third order.(3)documentation (e.g. sample size, test method, statistical analysis, build direction, post processing) for SLM generated specimens might be helpful to ensure comparability between conducted studies.Experiments on the fatigue behavior of 1.4542 stainless steel as-built and machined specimens as well as of thin-walled as-built actuator housings were performed.

A uniaxial test setup with a stress-ratio of R = 0 (pulsating tension) and test runs up to N = 107 cycles was selected. From the results the following conclusions can be drawn:The fatigue strength at N = 107 cycles of as-built specimens (σmax = 219 MPa, σa = 110 MPa) is significantly lower than of machined specimens (σmax = 492 MPa, σa = 246 MPa) due to a high surface roughness. The experimentally determined surface roughness reduction factor γ is 0.44. - Crack initiation starts for as-built and machined specimens from the surface. Therefore, manufacturing defects in the inside of the SLM generated specimens are not the reason for fatigue failure.Machined specimens show a higher sensitivity towards stress in the finite life region compared to as- built specimens.

This relation is shown through the slopes of the S-N curves in the relevant cycle regime.The endurable peak stress of thin-walled as-built actuator housings at 107 cycles is significantly above the fatigue strength of as-built specimens (factor 1.57 to 1.94). An explanation can be found in the stress gradient within the thin-walls due to the geometry. Stress gradients χ above 3.7 indicate that only a small volume fraction is exposed to high stresses; on average, the stress is significantly below the peak stress. S-N curves support the stress gradient concept that with increasing stress gradient the fatigue strength increases.The thin-walled actuator housings demonstrate fatigue strength at N = 107 cycles under application relevant loads. In addition, vertical and horizontal scattering of the S-N curves are below the values of the specimens.

In addition, vertical and horizontal scattering of the S-N curves are below the values of the specimens. This allows an approach of fatigue life with a low safety factor.Future research should be extended to thin-walled housings with other design parameters to prove the stress gradient concept. It is also desired to obtain valid assessment models for fatigue behavior of SLM generated parts without conducting extensive test series.AcknowledgmentsThis work was funded by the LOEWE Research Center AdRIA “Adaptronic Research, Innovation, Application”, grant number III L 4 – 518/14.0004 (2008). The authors would like to thank the Hessen State Ministry of Higher Education, Research and the Arts for the funding.ReferencesAdib-Ramezani, H. and Jeong, J. (2007), “Advanced volumetric method for fatigue life prediction using stressgradient effects at notch roots”, Computational Materials Science, Vol. 39 No. 3, pp. 649–663.Arola, D. and Williams, C.

viewed PaperFATIGUE ANALYSIS OF SHORT CARBON FIBER REINFORCED COMPOSITE COMPONENTS MANUFACTURED USING FIBER-REINFORCED ADDITIVE MANUFACTURINGMithila Rajeshirke*, Ismail Fidan†, Ankit Gupta†, and Kari Mäntyjärvi‡Department of Mechanical Engineering, Tennessee Tech University, Cookeville, TN 38505, USA†Department of Manufacturing and Engineering Technology, Tennessee Tech University, Cookeville, TN 38505, USA‡Department of Future Manufacturing Technologies, University of Oulu, Kerttu Saalasti Institute, Pajatie 5, FI-85500 Nivala, FinlandAbstractFiber-reinforced additive manufacturing (FRAM) has become quite popular in several industries. The technology offers an opportunity to improve the existing mechanical performance of the part. This research study has presented a successful methodology to fabricate the FRAM- based composite parts with improved fatigue properties.

This research study has presented a successful methodology to fabricate the FRAM- based composite parts with improved fatigue properties. Most engineering applications are subjected to cycling loading which makes the fatigue study an important analysis. The scope of this paper is to present the fatigue properties of short carbon fiber-reinforced Polyethylene Terephthalate Glycol (SCFs/PETG) of 13.78% by weight. The fatigue behavior was analyzed by varying the 3D printing process parameters i.e., infill orientation (0°, 45°, and 90°), and infill layer heights (0.2 and 0.3 mm). The tests are carried out on 1600 N as a maximum load of fatigue cycle with a 0.1 stress ratio, for the specimens with 90° and 45° orientations with 0.2 and 0.3 mm layer heights. For 0° orientation, both 0.2 and 0.3 mm layer height specimens are applied to 2600 N as maximum load, keeping the stress ratio the same as 0.1.

For 0° orientation, both 0.2 and 0.3 mm layer height specimens are applied to 2600 N as maximum load, keeping the stress ratio the same as 0.1. Analysis of Variance (ANOVA) is used to statistically analyze the testing data to understand the influence of input variables on fatigue properties.Keywords: FRAM, Tensile strength, Fatigue strength, ANOVA, Fatigue lifeIntroductionThe well-known additive manufacturing (AM) technologies are Fused Deposition Modeling (FDM) also known as Fused Filament Fabrication (FFF), Selective Laser Melting (SLM), Stereolithography Apparatus (SLA), and Selective Laser Sintering (SLS). The FRAM is the additive manufacturing technology used to manufacture the components with fiber-reinforced composite materials, which integrates the advantages of AM and composite materials. Fiber- reinforced composite materials are the materials produced by using fiber as reinforcement in a matrix material.

Fiber- reinforced composite materials are the materials produced by using fiber as reinforcement in a matrix material. Commonly used polymer matrix materials are nylon, polyester, epoxy,149polycarbonate (PC), acrylonitrile butadiene styrene (ABS), polylactide (PLA), polyamide (PA), etc. The reinforcement elements have categories as material-based and structural-based. Material- based reinforcements include Kevlar, Carbon, Fiber Glass, etc. while, structural-based reinforcements incorporate short fiber, continuous fiber, unidirectional, randomly oriented fibers, etc. [1]. By appropriate selection of fiber, and fiber volume fraction, length, type, and orientation, the mechanical, electrical, and thermal properties can be improved [1]–[3]. Fig. 1 shows the schematic view for the FRAM process using SCF reinforced filament.Recent advancements in AM made it applicable for using polymeric structures in various structural and load-bearing applications.

Cyclic loads cause fatigue, which is the development of structural damage, which leads to catastrophic failure at lower stress if compared with normal mechanical loading [4]. So, it is of utmost importance to study the fatigue behavior of the composite parts manufactured by AM. FRAM allows the researchers to explore more options in the materials with a combination of various matrix materials, fiber materials, and variations in the form of fibers. Fatigue tests are conducted in various loading conditions such as tension, compression, bending, torsion, fracture mechanical, or a combination of these in case of more complicated fatigue analyses. In fatigue testing, the simplest stress sequence uses a constant stress amplitude in which all the cycles are identical as given in fig. 2. For each cycle, 𝜎, 𝜎m, 𝜎max, and 𝜎min are the alternating stress, mean stress, maximum stress, and minimum stress respectively. The stress ratio is expressed as, R =σminσmax.Fig. 1.

2. For each cycle, 𝜎, 𝜎m, 𝜎max, and 𝜎min are the alternating stress, mean stress, maximum stress, and minimum stress respectively. The stress ratio is expressed as, R =σminσmax.Fig. 1. Schematic view of FRAM process [5]150Fig. 2. Nomenclature of the testing parameters in constant amplitude loading [6]In the FFF process, first, the required geometry is designed using CAD software and then sliced into the layers accordingly. Various printing parameters need to be taken into account such as printing speed, plane, orientation, infill density, infill patterns, layer height, layer width, extrusion temperature, bed temperature, etc. [4]. Various studies showed that these parameters affect the printing quality and mechanical and thermal properties. Researchers have also reported the fatigue behavior of various materials by FFF. The following section discusses various research studies reported till now considering printing parameters and materials. Alberto D. Pertuz et. al.

The following section discusses various research studies reported till now considering printing parameters and materials. Alberto D. Pertuz et. al. [7] studied the continuous fiber-reinforced composites (CFRC), in which nylon is used as matrix material, under static as well as fatigue loading. The study has shown the effects of infill pattern, infill density, the orientation of the fibers, and the material of fibers (fiberglass, Kevlar, and carbon fiber). From the study, in the end, it is concluded that nylon matrix reinforced with carbon fiber shows better performance for fatigue response at 0° orientation. Astrit Imeri et. al. [9], [10] in their study have shown the effect of fiber orientation and fiber materials (carbon, glass, and Kevlar). The study concluded that carbon fiber with zero and one concentric ring has high resistance to failure. Furthermore, Mst Faujiya Afrose et. al.

The study concluded that carbon fiber with zero and one concentric ring has high resistance to failure. Furthermore, Mst Faujiya Afrose et. al. [11] have investigated the effects of build orientations on the fatigue behavior of FDM processed-PLA under 50%, 60%, 70%, and 80% of ultimate tensile strength of respectively oriented specimens. The study concluded that specimens built in 45° achieved the highest fatigue life compared to specimens built in X and Y orientation. J. Antonio Travieso-Rodriguez et. al. [8] in their study explained the fatigue behavior of a polylactic acid-based composite reinforced with wood fabricated by FFF through rotating bending fatigue tests. The effects of layer height, nozzle diameter, infill density, and extrusion velocity are verified in this study, from which except extrusion velocity, all other parameters are influential on the fatigue life.

A thorough search of relevant literature yields that there is a limited number of studies that have been reported dealing with the fatigue behavior of SCFs/PETG with different printing parameters.The common advantages of the PETG material are, that it provides high processability, excellent chemical resistance, good tensile toughness, flexibility, high durability, low shrinkage,151and good adhesive properties between the layers [12]–[14]. The right amount of reinforcement of short carbon fiber in the matrix material has shown better mechanical properties [5]. Considering the requirement of industrial applications, it is highly necessary to study more and more options of commercially available materials under fatigue loading. Furthermore, a comprehensive understanding of the effects of process parameters on the fatigue behavior of FRAM printed parts is necessary for its further application.

Furthermore, a comprehensive understanding of the effects of process parameters on the fatigue behavior of FRAM printed parts is necessary for its further application. In this study, the FRAM is used to fabricate the specimens made by SCFs/PETG is the study aims to present the fatigue behavior of the chosen material taking into account the printing parameters such as layer height and infill orientation in the XY plane. The ANOVA is used to analyze the data obtained from the experimental results.

The ANOVA is used to analyze the data obtained from the experimental results. From the literature review, it is concluded that no in-depth investigation has been done to show the effect of layer height and printing orientation on SCFs/PETG.Materials and MethodsIn this section, the method of the investigation of the process parameters on the fatigueproperties of FRAM printed parts by SCFs/PETG is presented.Phase I: Set up of FRAM and DoEMaterial, geometry, and machineThe filament used for the research is commercially available SCFs/PETG (or carbon fiber reinforced PETG (CF-PETG)) by Push Plastic. As discussed earlier, PETG is a strong, versatile material having high heat resistance. Unlike some of the other materials, it does not have warping issues. PETG can also be used for food containers and tools used for food consumption. Reinforcement of carbon fibers improves the properties of the matrix material such as high tensile strength, high stiffness, and low weight [15].

Reinforcement of carbon fibers improves the properties of the matrix material such as high tensile strength, high stiffness, and low weight [15]. As compared to continuous fibers, SCFs can be molded into complex shapes. SCFs are easy to manufacture and low cost. The properties from the manufacturer’s data sheet are as given in Table 1.Table 1. Material Properties of SCFs/PETG [16] Values 14,000 (PSI) or 96.52 (N/mm2) 4.00% 20,500 (PSI) or 141.34 (N/mm2) 1,000,000 (PSI) or 6894.75(N/mm2) 1.31Test Method ASTM D638 Tensile Elongation at Break ASTM D638 ASTM D790 ASTM D790 ASTM D792Properties Tensile Strength at YieldFlexural Strength Flexural Modulus Specific GravityTo determine the percentage by weight of carbon fibers in the matrix material, samples or filaments are cut out and incinerated in thermogravimetric analysis (TGA) system i.e., SDT 650 to burn off the PETG.

Using a heating program in the presence of Nitrogen the TGA is performed with a temperature profile ranging from 25 °C to 786.94 °C with a rate of 10 °C per minute. The obtained result concludes that the CF is 13.78% by weight. The graphical representation of the TGA is shown in fig. 3.The geometry selected is based on the ISO 11782-1:1998(E) [17] as given in fig. 4. The SOLIDWORKS is used to generate the CAD model and then saved in STL format. This STL file152is imported into the PRUSA slicing software where the printing parameters are assigned as given in table 2. The printing plane is XY and the infill orientations used are 0°,45°, and 90° as shown in fig. 5(a) and 5(b). After slicing the G-code is generated and it is provided to the PRUSA i3 MK3 printer with a hardened steel nozzle of 0.4 mm diameter.TGA Analysis120100)%80(t h g i e60W402013.78%00100200300400500600700800900Temperature (°C)Fig. 3.

Levels of factors for DoE LH (mm) Levels 0.2 1 0.3 2 - 3O 0° 45° 90°Phase II: ExperimentBefore initiating the fatigue tests, it is required to have preliminary tensile test data to define the fatigue maximum and minimum loading conditions. The maximum load of the fatigue cycle should not increase more than the ultimate tensile load. From the literature review, tensile strength is the lowest for 90° if compared to 45° and 0° infill orientation[11], [18]. Mithila Rajeshirke et. al [18] in their study have reported the lowest tensile load for CF-PETG with 90° infill orientation as 1909.48 N as compared to 45° (2158.05 N) and 0° (2574.35 N) infill orientations.154The same specimen geometry is used to determine the ultimate tensile load sustained by the specimens by 90° infill orientation with 0.2 mm LH.

A total of three replicates are tested under static tensile testing with a strain rate of 5 mm/min to determine the ultimate tensile load and then the average of these three tested samples is determined. The results obtained for the tensile test are shown in Table 4. Ultimate tensile strength is obtained by dividing the area of the broken specimen. 810E4-15 Dynamic Test System is used for the tensile and fatigue tests. It has a 15 kN axial load capacity for static and fatigue testing applications.Specimen No.1 2 3Table 4.

810E4-15 Dynamic Test System is used for the tensile and fatigue tests. It has a 15 kN axial load capacity for static and fatigue testing applications.Specimen No.1 2 3Table 4. Result for Tensile Test (LH - 0.2, O - 90°) Ultimate Ultimate Tensile Stress Tensile Load (N/mm2) (UTL) (N) 39.85 2010.69 41.19 2079.93 39.29 1996.46Average of UTL (N)2029.0280% of the Average UTL (N)1623.22Considering the approximate value of 80% of that load i.e., 1600 N is applied as the maximum load and 160 N for the minimum load for the fatigue cycle to the specimens with orientations 90°, 45° and layer heights 0.2, 0.3 mm for fatigue testing, which eventually implicates the stress ratio as 0.1. The frequency used is 3 Hz. All the tests are carried out at 53-60% of humidity and 68-70 °F temperature. Newton test machine controller/software is used to control the test parameter. Proportional/Integral/Derivative (PID) values need to be set up to control the waveform accurately.

Newton test machine controller/software is used to control the test parameter. Proportional/Integral/Derivative (PID) values need to be set up to control the waveform accurately. PID values used are 1400/1200, 100, and 0 respectively. The test is load controlled fatigue test, as loads are controlled. The response to the tests is recorded in the form of a number of cycles.When the same loads are applied to the specimens (two) of orientation 0° and layer height 0.2 mm, the number of cycles under fatigue load obtained exceeded the set limit i.e., 50,000. Since the time required for such long cycles is exorbitant, it is decided to perform the tensile tests on the same specimen and then fatigue tests considering the approximate value of 80% average ultimate load sustained by the specimen as maximum load in the fatigue cycle. The results obtained for the tensile tests are shown in Table 5.Specimen No.1 2 3Table 5.

The results obtained for the tensile tests are shown in Table 5.Specimen No.1 2 3Table 5. Result for Tensile Test (LH - 0.2, O - 0°) Ultimate Ultimate Tensile Stress Tensile Load (N/mm2) (UTL) (N) 58.03 3160.51 63.82 3409.73 66.30 3398.20Average of UTL (N)3322.8180% of the Average UTL (N)2658.24Considering the approximate value of 80% of that load i.e., 2600 N is applied as the maximum load and 260 N for the minimum for fatigue cycle load to the specimens with layer heights 0.2, 0.3 mm for fatigue testing. The stress ratio and frequency considered are 0.1 and 3 Hz respectively.155Result and DiscussionIn this section, the results obtained from the fatigue tests and statistical analysis arediscussed.Phase I: ResultsThe result obtained from the fatigue tests is shown in Table 6. The response is mentioned in terms of the number of cycles sustained by the specimens. A total of 12 replicates are tested for each factor level of DoE.

The response is mentioned in terms of the number of cycles sustained by the specimens. A total of 12 replicates are tested for each factor level of DoE. The load applied to each of these specimens is discussed in the experiment section. The specimens after testing for each orientation and layer height are as shown in fig. 6.Table 6. Results of Fatigue Test (Response – Number of cycles)O LH / Replicates 1 2 3 4 5 6 7 8 9 10 11 120.23260 5948 5163 5604 9169 10302 2354 4782 4301 6090 9015 867790°0.35106 7383 4393 4031 5387 4581 5512 5677 6122 9470 12252 59660.23294 5300 6464 6708 3711 7192 4024 4228 3766 7058 6026 1019445°0.38392 8168 9398 5339 6540 4716 7356 6912 8246 6606 8344 104240.2362 644 176 888 504 816 674 554 696 262 660 5610°0.330 140 56 23 82 33 42 106 31 35 53 106Fig. 6.

6. The specimens after fatigue tests for the orientations and layer height as – a) 90°,0.2mm, b) 45°, 0.2 mm, c) 0°, 0.2 mm, d) 90°, 0.3mm, e) 45°, 0.3mm, and f) 0°, 0.3 mm156Phase II: Statistical AnalysisFor O – 90° and 45°With the aid of SAS (Statistical Analysis System) software, the ANOVA analysis is performed on the fatigue results data for 90° and 45° orientations with 0.2 and 0.3 mm (first four columns of table 6), obtained from the testing. The reason to consider only these orientations (90° and 45°) for ANOVA is, that these specimens were applied to the same loading i.e., 1600 N as a maximum load, unlike 2600 N load was applied as the maximum load to 0° orientation specimens. The analysis is done to investigate the individual input parameters and their interactions with the fatigue behavior. The ANOVA model chosen is shown in eq.

The analysis is done to investigate the individual input parameters and their interactions with the fatigue behavior. The ANOVA model chosen is shown in eq. (1):Yij = μ + αi + βj + (αβ)ij + εij (1)Where μ is a constant (overall mean), αi is the main effect for factor A (Layer Height) at the ith level, βj is the main effect for factor B (Orientation) at the jth level, (αβ)ij is the interaction factor, and εij is the error term. The probability value (p-value) aids to determine whether the input variable is significant or not, by comparing it with the alpha value of 0.05. If the p-value is smaller than 0.05, the input variable is considered to be significant. The ANOVA table and interaction plot are shown in Table 7 and fig. 7 respectively.Table 7.

If the p-value is smaller than 0.05, the input variable is considered to be significant. The ANOVA table and interaction plot are shown in Table 7 and fig. 7 respectively.Table 7. ANOVA of fatigue results for main effects and their interaction with p values Sum of squares Mean square F-value p-value SourceDegree of Freedom 1 1 1LH O LH*O11908872.71 1024221.78 9643867.8911908872.71 1024221.78 9643867.892.66 0.23 2.150.1102 0.6348 0.1494Fig. 7. The interaction plot for the response (number of cycles) for 90° and 45° infillorientations with 0.2- and 0.3-mm layer heights157From the statistical analysis, as p values for LH, O, and their interaction (LH*O) all are above 0.05, meaning that the evidence suggests there is no significance either for main effect factors (O and LH) or their interactions. From the interaction plot, fig. 7, it can be observed that the response (i.e.

From the interaction plot, fig. 7, it can be observed that the response (i.e. number of cycles) for 90° and 0.2 mm LH varies from 2354 to 10302, and for 0.3 mm LH from 4031 to 12252. For 45° infill orientation with 0.2 mm LH it varies from 3711 to 10194 and with 0.3 mm LH, from 4716 to 10424. This means, that for the lower load such as 1600N, neither the layer heights nor orientations make any significant difference in the fatigue behavior if the values for layer heights are 0.2 and 0.3 mm and orientations are 90° and 45°. The large variation in the data is because of the fact that printed parts have unavoidable defects such as pores, voids, and poor adhesion which may cause the parts to fail at a very low number of cycles. Also, manual errors while testing the specimens such as minor variations in adjusting the grippers may affect the loads and eventually the fatigue behavior.

Also, manual errors while testing the specimens such as minor variations in adjusting the grippers may affect the loads and eventually the fatigue behavior. As 90° and 45° both infill orientations are not in the direction of the applied loads, are giving very low responses, and do not show much significant difference even with respect to the layer heights.For O – 0°As mentioned in the experiment section, when 1600 N as the maximum load is applied to the specimens (two) of orientation 0° and layer height 0.2 mm, the number of cycles under fatigue load obtained exceeded the set limit i.e., 50,000, as the fibers printed in the direction of loading conditions makes it more resistant to the applied load.The obtained test results for fatigue behavior for 0° orientation and the load of 2600 N as maximum load, are presented in the Box-Plot in fig. 8. Further details about applied loading conditions are discussed in the experiment section.

8. Further details about applied loading conditions are discussed in the experiment section. From the chart, it is concluded that 0.2 mm layer height shows better results than 0.3 mm. In the case of smaller layer heights results in better conditions for melting each subsequent layer during the contact of the hot extrudate with the previous print layer. This phenomenon may have a decisive influence on forming the interlayer adhesion.Fig. 8. The fatigue results in chart form for O - 0°158ConclusionUsing FRAM technology, the specimens were fabricated with SCFs/PETG. TGA is used to determine the weight percentage of SCF in SCFs/PETG composite. Tensile tests are performed to determine the loading conditions required for fatigue testing. The fatigue tests are performed with respect to various printing parameters and are reported in this article.

The fatigue tests are performed with respect to various printing parameters and are reported in this article. Based on the analysis of the data obtained from the fatigue tests following conclusions have been made.For the low loads such as 1600N (as maximum load), 90° and 45° orientations do not show any significant difference in the fatigue behavior. Even for the different layer heights i.e., 0.2 and 0.3mm there is no significant difference. From the obtained result data, there is a large variation in the response. The number of cycles of more than 10,000 could be the result of lower gripping pressure. Gripping the specimens was done manually.

The number of cycles of more than 10,000 could be the result of lower gripping pressure. Gripping the specimens was done manually. While a very low number of cycles could have resulted from unavoidable defects in printing such as voids, poor adhesion, etc.For the low loads as mentioned above (1600N as maximum load), the 0° orientation can sustain more than 50,000 cycles.For the higher loads such as 2600N (as maximum load), the 0° orientation with 0.2mm layer height shows better performance than the 0.3mm layer height. In the case of smaller layer heights results in better conditions for melting each subsequent layer during the contact of the hot extrudate with the previous print layer. This phenomenon may have a decisive influence on forming the interlayer adhesion. Future ScopeIt should be mentioned that limited printing parameters are studied for this research and the importance and influence of various other parameters may be different.

Future ScopeIt should be mentioned that limited printing parameters are studied for this research and the importance and influence of various other parameters may be different. Therefore, future research should focus on more parameters such as nozzle diameter, stacking sequence, and carbon fiber percentage. Furthermore, the microstructural analysis should be carried out to have in-depth knowledge about the failure of fibers and matrix. To determine the fatigue strength of the material should also be taken into consideration. The reproducibility of FRAM in terms of fatigue properties should be verified on different geometry as well. Then more universal guidance can be accepted for fatigue behavior.159AcknowledgementsThis material is based upon work supported by the National Science Foundation under Grant No. 1902437, “Mobile Additive Manufacturing Platform for 21st Century STEM Workforce Enhancement”.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFATIGUE BEHAVIOR AND FAILURE MECHANISMS OF DIRECT LASER DEPOSITED INCONEL 718Alexander S. Johnson1, Shao Shuai1, Nima Shamsaei2,*, Scott M. Thompson2, Linkan Bian31Center for Advanced Vehicular Systems (CAVS), Mississippi State University, MS 397622Department of Mechanical Engineering, Auburn University, Auburn, AL 368493Department of Industrial and Systems Engineering, Mississippi State University, MS 39762Corresponding author: shamsaei@auburn.eduAbstractInconel 718 is considered as a superalloy with a series of superior properties such as high strength, creep-resistance, and corrosion-resistance.

Additive manufacturing (AM) is particularly appealing to Inconel 718 because of its near-net-shape production capability to deal with the poor machinability of the alloy. However, AM parts are prone to porosity which is detrimental to the alloy’s fatigue properties. As such, further understanding of the fatigue behavior of AM Inconel 718 is much needed. The room temperature fatigue behavior of AM Inconel 718 produced by an Optomec Laser Engineered Net Shaping (LENSTM) system is investigated in this study. Build conditions are carefully controlled to minimize the scatter in fatigue data. The specimens are tested after being subject to a standard heat treatment. Fully reversed strain controlled fatigue tests are performed on round specimens with straight gage section at strain ranges of 0.1% to 1%. Fracture surfaces of fatigue specimens are inspected using a scanning electron microscope.

Fracture surfaces of fatigue specimens are inspected using a scanning electron microscope. Results are compared to literature for both AM and wrought materials.Keywords: Fatigue; Failure Mechanisms; Tensile Behavior; Microstructure; AdditiveManufacturing; Superalloys1. IntroductionInconel 718 is a precipitation hardened nickel-based super alloy used mainly in extreme temperature aerospace applications, such as for turbine blades in jet engines, liquid-fueled rocket components, and cryogenic containers [1]. Inconel alloys are well known for their toughness and difficulty to machine [1]. To circumvent a large portion of the traditional machining time, additive manufacturing (AM) can be used to produce near net-shape parts with only light machining necessary to produce the final product. In general, Fatigue accounts for ~90% of material failures, however, to date, most fatigue research on Inconel 718 has centered on traditionally manufactured Inconel specimens [2-9].

In general, Fatigue accounts for ~90% of material failures, however, to date, most fatigue research on Inconel 718 has centered on traditionally manufactured Inconel specimens [2-9]. In general, it has been found that fatigue behavior of wrought Inconel 718 begins with a short period of cyclic hardening followed by cyclic softening for the remainder of its fatigue life [6]. In both the low-cycle and high-cycle fatigue regimes, cracks typically initiate499from the surface of specimens and grow to fracture [2, 8]. Most crack growth has been found to be of a transgranular, faceted mode [2-5, 9].Strain controlled fatigue data is important in that they can provide insights into both the elastic and plastic material behavior during fatigue loading [10]. Very little strain controlled data has been published for wrought Inconel 718 [9], with no strain controlled fatigue data available for AM Inconel 718 to the authors’ best knowledge.

Very little strain controlled data has been published for wrought Inconel 718 [9], with no strain controlled fatigue data available for AM Inconel 718 to the authors’ best knowledge. It is of interest to assess the applicability of AM methods to Inconel 718 and compare its fatigue performance to its traditionally manufactured form. If the performance of the AM material is comparable to wrought, then AM can more justifiably be used to quickly replace broken parts, or more easily manufacture Inconel 718 parts in industry. In this study, Laser Engineered Net ShapingTM (LENS), a blown-powder/laser-based Directed Energy Deposition AM method, is employed for generating Inconel 718 fatigue specimens. Specimens are tested using strain control in order to investigate the fatigue behavior of AM Inconel 718 in both short and long life regimes. The effects of the heat treatment procedure on the microstructure of the material is examined.

The effects of the heat treatment procedure on the microstructure of the material is examined. The strain-life data is compared with the wrought material under the same heat treatment conditions. The specimen fracture surfaces are inspected to understand fatigue crack initiation and propagation characteristics and to compare to reported results in the literature.2. Experimental ProcedureDuring the LENS process, multiple nozzles are used to inject powder metal into a centrallaser beam creating a melt pool on a substrate of similar material to the powder atop a CNC stage. The stage is then moved in the x/y plane (horizontal) relative to the laser beam, in a user-defined path, in order to deposit a line of molten metal. The planar sum of these lines, or tracks, form layers while multiple, stacked layers then form a three dimensional (3D) part that is based on a user- supplied 3D CAD model.

The planar sum of these lines, or tracks, form layers while multiple, stacked layers then form a three dimensional (3D) part that is based on a user- supplied 3D CAD model. During this process, inert gas is pumped into the process chamber to reduce material oxidization.An argon-purged OPTOMEC LENSTM 750 system equipped with a 1 kW Nd:YAG laser is employed for specimens fabrication. In order to ensure dense, near net-shape parts, a series of preliminary line tests are performed to best determine an operable set of process parameters, i.e.: laser output power, scanning speed and powder flow rate. Using various scan speeds, six lines of arbitrary length are deposited adjacently at different combinations of powder flow rates and laser powers. These deposited lines are cut from the substrate and subsequently sectioned to inspect their cross-sectional shape and fusion with the substrate.

These deposited lines are cut from the substrate and subsequently sectioned to inspect their cross-sectional shape and fusion with the substrate. Using satisfactory process parameters, 21 cylindrical rods were manufactured using a method consisting of a circular contour deposit with slightly overlapping linear deposits used inside the contour to create the interior of each layer. The direction of the linear deposits are rotated 90 degrees in the x/y plane every layer. Parameters are kept constant throughout the build process. All specimens were printed on the same wrought Inconel 718 substrate spaced approximately 75 mm apart. Previous builds were allowed to cool for several minutes before the next build is started. Specimens were removed from the plate with a hacksaw.500All of the builds were machined into round tensile and fatigue specimens according to ASTM Standard E606 [11] for testing, except one which remained un-machined for inspection of its as-built microstructure.

Specimens’ geometry is shown in Fig. 1. Fatigue tests were conducted at room temperature in ambient lab atmosphere. All specimens were subject to the standard heat treatment for Inconel 718; 940°C for two hours followed by out-of-furnace air cooling, aging at 718°C for eight hours followed by cooling at a rate of 56°C/hr to 621°C and held for another eight hours followed by air cooling [9].Figure 1. Additive manufactured Inconel 718 fatigue specimens per ASTM E606 [11] (dimensions in mm).Both a heat treated and as-built specimens were each sectioned, mounted, and polished for microstructure analysis. Etching was accomplished using Waterless Kalling’s etchant. Three specimens were monotonically tested to obtain tensile properties of the material. Fully reversed fatigue tests were run on a MTS Landmark system under strain controlled condition at fully reversed strain amplitudes ranging from 0.1% to 1% with at least two tests at most strain levels.

Before testing, the gage sections of all specimens were polished to a mirror finish along the loading direction to remove circumferential surface machining marks. After a number of cycles, some tests at high cycle regime were switched to load control for the remainder of their fatigue lives as was also done in the work of Brinkman and Korth [9]. After failure, fracture surfaces were cut from the fatigue specimens and examined via scanning electron microscopy (SEM).3. Experimental ResultsMechanical properties of the heat treated AM specimens measured in this study and reported elsewhere in the literature, as determined through monotonic testing, are provided in Table 1. If multiple tests of the same type were performed, values were averaged.

If multiple tests of the same type were performed, values were averaged. The tensile mechanical properties of the investigated AM specimens demonstrate to have similar yield stresses (standard deviation of 4.2%), ultimate tensile stresses (UTS) (standard deviation of 1.4%), and elongation to failure (standard deviation of 2.3%) Compared to the wrought Inconel 718 data presented in Table 1, the AM Inconel in this study has slightly lower yield stress and higher UTS.501Table 1. Tensile Properties of Heat Treated Inconel 718.0.2% Yield Stress (MPa) UTS (MPa) Elongation (%)This study Blackwell [12] Zhao et al. [13] Amsterdam & Kool [14] Wrought [12, 13]895 1257 1133 891 11131514 1436 1240 1213 135319.7 13.0 9.0 10.4 16.0Microstructure of the as-built specimen and a specimen after heat treatment are presented in Fig. 2. Builds are sectioned into segments consisting of the grip section and gage section.

2. Builds are sectioned into segments consisting of the grip section and gage section. These sections are cut in half longitudinally (Sections B and C) to view the inner sections of the build. For the heat treated sections, a circular cross section of the grip section is also examined (Section: A). Microstructure is revealed using Waterless Kalling’s solution.Lack of fusion between layers is prevalent in the first several layers deposited near the substrate (Fig. 2(a)). However, these type of defects become nearly non-existent, with only general spherical porosity present in layers further from the substrate (Fig. 2(b)), i.e. in the gage section of specimens. The lack of fusion is confined to the observed grip sections of the selected specimens. Microstructures of specimen regions located near the substrate are found to be mainly comprised of a dendritic structure even after heat treatment (Fig. 2(c)).

Microstructures of specimen regions located near the substrate are found to be mainly comprised of a dendritic structure even after heat treatment (Fig. 2(c)). Further from the substrate after heat treatment, the dendritic structure is found to have mostly dissolved into the general grain structure, though there are residual laves and δ particles present from the dendritic structure (Fig. 2(d), circled) as discussed by Qi et al [15]. Small portions of the dendritic structures remained in the gage section after heat treatment (Fig. 2(e), circled). Average grain size of the heat treated gage section is found to be 43 μm. Figure 2(f) shows the microstructure of the gage section of the as- built specimen. In comparison to the microstructure seen in Fig. 2(e), the grains are found to be much larger and elongated (average of ~170 μm in size along the elongated direction), with more variation in grain size with faint dendritic structure within the grains.

Based on microstructural observations, it is clear that the employed heat treatment created more similarly sized grains as well as made them smaller.Fatigue data, as well as the Coffin-Manson curve fit, are provided in a strain versus life plot, as shown in Fig. 3. It may be seen that the collected data follows a smooth, consistent curve with little scatter. Equation (1) details the specifics of the Coffin-Manson fit for the elastic and plastic portions of the AM Inconel 718 fatigue-life curve [16], i.e.:∆𝜀 2=′ 𝜎f 𝐸(2𝑁f)𝑏 + 𝜀f′(2𝑁f)𝑐 (1)∆𝜀′ is the fatigue strength coefficient, 𝐸 is the ′ is the (2𝑁f)𝑏 expressesIn Eq. (1),is the total strain amplitude, 𝜎f2modulus of elasticity, 2𝑁f is the reversals to failure, 𝑏 is the fatigue strength exponent, 𝜀f fatigue ductility coefficient, and 𝑐 is the fatigue ductility exponent. The term′ 𝜎f 𝐸502′(2𝑁f)𝑐 expresses the plastic portion.

For comparable strain amplitudes, the fatigue lives of the wrought material were two to three times as long as the AM specimens.To more directly compare experimental results herein to those of Amsterdam and Kool [14], whom used a different loading ratio during their fatigue testing, the stress values based on each strain interval are determined using measured forces and gage section cross-sectional area. The R-value for the data herein is calculated for each specimen and averaged for all tests. An equivalent stress versus life plot is generated and the Smith-Watson-Topper mean stress correction is employed, as shown in Eq. (3) [17], i.e.:𝜎eq = 𝜎max√1 − 𝑅 2(3)where 𝜎eq is the equivalent stress amplitude for a test in fully reversed conditions, 𝜎max is the maximum stress for the test data being normalized, and R is the load ratio for the test being normalized. The data correlates very well between this study and the work of Amsterdam and Kool, as shown in Fig. 4.Figure 3.

The data correlates very well between this study and the work of Amsterdam and Kool, as shown in Fig. 4.Figure 3. Strain-life fatigue behavior of AM Inconel 718 and its comparison to wrought Inconel 718 [9]. Arrows indicate runout specimens, and the circled point is a specimen that failed much earlier than expected.504Figure 4. Comparison of normalized R-value stress-life fatigue data of AM Inconel 718 generated in this study with literature [14].Examples of fracture surfaces obtained from fatigue tests in high and low cycle regimes are presented in Figs. 5 and 6, respectively. All specimens exhibited a distinct crack growth regions (Fig. 5(a) and (c), left side of the red dashed line) and final fracture regions (Fig. 5(a) and (c), right side of the dashed line). Crack growth regions are characterized by fine striations (circled in Fig. 5(b)) running perpendicular to the crack growth direction (white dashed arrows in Figs. 5 and 6).

Crack growth regions are characterized by fine striations (circled in Fig. 5(b)) running perpendicular to the crack growth direction (white dashed arrows in Figs. 5 and 6). A magnified view of the boundary between the regions on the corresponding surface to Fig. 5(a) is provided in Fig. 5(c). Note that in Fig. 5(c), the parallel line features on the upper right quadrant are part of the final fracture region and are aligned dimples instead of striations. The final fracture regions are distinguished by the high amount of dimpling on the fracture surface (Fig. 5(c), right side of the line), indicating the final fracture is ductile in nature. Most of the crack growth regions are generally flat and of transgranular mode with clear river marks running back towards the initiation point (Fig. 5(d), red solid arrows).In the specimens shown in Figs. 5 and 6, striations are present on flatter cleavage surfaces as circled in Fig. 5(b) and boxed in Fig. 6(b).

5(d), red solid arrows).In the specimens shown in Figs. 5 and 6, striations are present on flatter cleavage surfaces as circled in Fig. 5(b) and boxed in Fig. 6(b). Many exposed pores can be seen on the high cycle and low cycle fatigue fracture surface in Fig. 5(d) and Fig. 6(d) respectively, as marked by red solid arrows. Pores this close to the fracture surface likely played a role in reducing the initiation and early growth of fatigue cracks. The fracture surface examined in Figs. 5(e) and 5(f) belong to the specimen from the circled data point in Fig. 3. This sample failed much earlier than other specimens at a strain amplitude of 0.2%. The lower fatigue life likely stems from initiation point in Fig 5(f) (indicated by the arrow), a pore exposed to the specimen surface.505Figure 5. Fatigue fracture surfaces of AM Inconel 718 from high-cycle fatigue tests. Crack propagation directions marked by white dashed arrows.

Fatigue fracture surfaces of AM Inconel 718 from high-cycle fatigue tests. Crack propagation directions marked by white dashed arrows. (a) & (b) 0.3% strain amplitude test with notable striations present in (b) (in circled area), exposed pores indicated by red solid arrows. (c) Transition from crack growth to final fracture (indicated by dashed arc). (d) River marks (indicated by red solid arrows) leading back towards initiation site (circled). (e) & (f) 0.2% strain amplitude test that failed earlier than other tests. Crack initiation point is indicated by red solid arrow in (f).4. DiscussionThe mechanical properties of investigated LENS Inconel 718 are similar to AM and wrought reported in the literature. Other data on LENS-type AM Inconel 718 comes from Blackwell [12] and Zhao et al. [13] is presented in Table 1 for comparison.

Other data on LENS-type AM Inconel 718 comes from Blackwell [12] and Zhao et al. [13] is presented in Table 1 for comparison. With the full heat treatment data presented by Blackwell, the UTS of the specimens manufactured in this study are similar, within 100 MPa of the average of the three specimens tested. However, the yield and elongation to failure differed by a greater degree. The yield stress reported by Blackwell was significantly greater than the specimens produced in this study. The elongation, on the other hand, was smaller in Blackwell’s report by an average of over 6% [12].506Figure 6. Fatigue fracture surfaces of AM Inconel 718 from low-cycle fatigue tests. Crack propagation directions marked by white dashed arrows. (a) & (b) 1% strain amplitude test, (b) Initiation of the crack (circled), cracks forming from the surface (indicated by red solid arrows). Striations near initiation region (boxed). (c) & (d) 0.6% strain amplitude test, (d) Crack initiation region.

Striations near initiation region (boxed). (c) & (d) 0.6% strain amplitude test, (d) Crack initiation region. Many exposed pores in the crack initiation region (several pores indicated by red solid arrows).For the data presented in the work of Zhao et al., the same relationship exists with the yield stresses in this study as it has with Blackwell’s data. The yield in the current study is lower to about the same degree, though the UTS and the elongation to failure also differ to a greater degree than seen in Blackwell. The specimens produced in this study are much more ductile as compared to Zhao et al. [13], with elongations to failure of 10% more and with an average UTS of 250 MPa more. It is possible that differences in the heat treatment process, such as the addition of hot isostatic pressing (HIPing) in Zhao et al. [13], were responsible for the differences in material behavior.

It is possible that differences in the heat treatment process, such as the addition of hot isostatic pressing (HIPing) in Zhao et al. [13], were responsible for the differences in material behavior. An average of the mechanical properties presented by Amsterdam and Kool [14] had nearly the same yield stress as the material investigated in this study. The average UTS and elongation to failure presented by Amsterdam and Kool were comparable to Zhao et al., and thus, have the same relationship to this study. As compared to averaged wrought Inconel 718 data assembled from Zhao et al. and Blackwell’s works, the mechanical properties of the specimens used in the current study have higher UTS and lower yield stress, though it has more comparable elongation to failure [12, 13].507Microstructural comparisons can be easily made with the abundance of literature on the microstructure of AM Inconel 718. Blackwell [12], Zhao et al. [13], Qi et al. [15], and Liu et al.

Blackwell [12], Zhao et al. [13], Qi et al. [15], and Liu et al. [19] all report on the heat treated and as-built microstructure with Chen et al. [18] reporting solely on as-built. The most common microstructural feature seen in this work, as well as reported in the literature is the significant presence of dendritic microstructure prior to heat treatment in AM Inconel 718 growing approximately in the build direction, as can be seen in the upper region of Fig. 2(b). The long, thin dendrite structures are most likely caused by a high cooling rate during the solidification of the deposited material [12, 13, 18, 19]. After heat treatment, most of the large, elongated grains present in the as-built specimen (Fig. 2(f)) form smaller and more equiaxed grains (Fig. 2(e)). These results are consistent with the results of Zhao et al. [13] and Liu et al. [19]. An interesting phenomenon is presented in Fig. 2(d), which is explained in the research presented by Qi et al. [15] in more detail.

[13] and Liu et al. [19]. An interesting phenomenon is presented in Fig. 2(d), which is explained in the research presented by Qi et al. [15] in more detail. The spotted pattern observed (Fig. 2(d), circled) in the microstructure after heat treatment are residual laves and δ phases left over from the dendritic structure present in the as-built material. The solution annealing temperature was not high enough to fully dissolve these phases back into the material matrix [15].Lack of fusion between the substrate and early deposited layers is also a problem that occurred in the “cross-bonded” specimens presented in Blackwell’s work [12]. In this study, lack of fusion in the first several layers of builds near the substrate was also present. These defects are most likely due to rapid cooling and instability of the melt pool as the non-preheated substrate absorbs and disperses directed energy during early layer deposits.

These defects are most likely due to rapid cooling and instability of the melt pool as the non-preheated substrate absorbs and disperses directed energy during early layer deposits. As the build gets further away from the substrate, the previously-deposited layers retain more heat and allow for better fusion of subsequent layers. Further investigation with a heated substrate would have to be performed to extrapolate on this build behavior. As large amounts of porosity and lack of fusion tend to occur in the layers near the base of the build (Fig. 2(a)), it is recommended that critical structural portions of builds be built in layers further away from the substrate. Pores found in the gage section of the heat treated specimens are generally circular in shape and did not exceed 45 μm in diameter. This diameter is within the range of pore sizes, 25 and 70 μm, observed by Zhao et al.

This diameter is within the range of pore sizes, 25 and 70 μm, observed by Zhao et al. [13].The fatigue performance of investigated specimens has been found to be comparable to the AM Inconel 718 performance reported elsewhere [14], as seen in Fig. 4, and less than that of wrought material [9], as seen in Fig. 3. The maximum stress-life fatigue data of AM Inconel 718 presented by Amsterdam and Kool [14] with R = 0.1 has been normalized to equivalent stress-life data using the Smith-Watson-Topper method detailed in Equation (2) and compared to this study in Fig. 4. Heat treatment for the material tested by Amsterdam and Kool was similar to the treatment used in this study with the exception of a higher annealing temperature and different aging times, however, the overall heat treatment time remained the same.

Excluding a couple of outliers and the run-out tests for both data sets, this method equates the data nearly exactly, with the bulk of data lying in line with the data presented herein. The fatigue data for wrought Inconel 718 reported by Brinkman and Korth [9] is compared with the current study results in Fig. 3. Note that the heat treatment process used in both studies were the same. While fatigue lives are comparable in short life regime, fatigue lives of LENS Inconel 718 are at least an order of508magnitude shorter than the ones of the wrought counterpart.

While fatigue lives are comparable in short life regime, fatigue lives of LENS Inconel 718 are at least an order of508magnitude shorter than the ones of the wrought counterpart. Porosity present in the AM specimens is the main source of the discrepancy between the wrought and additive material behavior as they serve as crack initiation sites and speed up the fatigue process.Although the exact point of crack initiation could not be located for a few of the fracture surfaces, indications of fatigue crack initiation from pores, as well as the pores accelerating the fatigue growth could be found in other specimens. For instance, there are unmelted particles near the surface that have initiated the crack (Fig. 6(b)) in several specimens. The fracture surface presented in Fig. 5(e) and 5(f) comes from a 0.2% strain amplitude specimen (Fig. 3, circled) that failed much earlier in its fatigue life than others of the same strain amplitude.

5(e) and 5(f) comes from a 0.2% strain amplitude specimen (Fig. 3, circled) that failed much earlier in its fatigue life than others of the same strain amplitude. There appears to be a pore exposed to the surface of the part that initiated failure and this may explain the much shorter fatigue life. A few of the surfaces also had pores very close to the surface that have initiated fracture like those seen in the initiation region of Fig. 6(c) magnified in Fig. 6(d) (top arrow). The amount of exposed pores in the area indicates it played some role in the early crack growth. Most of the porosity present on the fracture surfaces near the outer surface of the gage section is less than 5 μm in diameter as seen in Fig. 6(d). Exposed pores are present in many of the fracture surfaces in the crack growth regions (Figs.

6(d). Exposed pores are present in many of the fracture surfaces in the crack growth regions (Figs. 5(d), 6(b), and 6(d)).For fracture surface comparisons, there exists only one study on the fatigue behavior of AM Inconel 718 in published literature to the best of authors’ knowledge. The work of Amsterdam and Kool [14] details the fracture surface appearance of AM Inconel 718. The description provided of the fracture behavior of the material matches the behavior observed in this study. Initiation from the surface of the material with final fracture regions indicated by the large amount of dimpling [14]. Similar ductile, dimpled final fracture behavior is demonstrated in wrought Inconel 718 studies performed by Ma et al. [3]. The crack propagation observed in specimens tested in this study are in a transgranular mode. This behavior is backed up by wrought crack growth behavior observed in literature [2-5, 9].5.

[3]. The crack propagation observed in specimens tested in this study are in a transgranular mode. This behavior is backed up by wrought crack growth behavior observed in literature [2-5, 9].5. ConclusionsThe mechanical properties, microstructure, fatigue life, and failure modes of additive manufactured (AM) Inconel 718 alloy has been investigated and the results have been discussed. AM Inconel 718 specimens were fabricated using the Laser Engineered Net Shaping (LENS) technique. All specimens were given a standard heat treatment and tested in the same manner as strain-controlled testing of wrought Inconel 718. Resulting fatigue behavior was compared with literature along with microstructure, mechanical properties, and fracture surfaces. The following conclusions can be drawn from acquired data:1. As-built AM specimens had elongated grains in the build direction as well as a large amount of dendritic structure.

The following conclusions can be drawn from acquired data:1. As-built AM specimens had elongated grains in the build direction as well as a large amount of dendritic structure. The issue of lack of fusion near the substrate during builds needs to be further explored in an attempt to produce denser builds. Preheating the substrate may alleviate this lack of fusion.2. Heat treatment of the specimens was found to create smaller and more uniform grain structure. Residual dendrites still existed in some portions of the investigated specimens.509In places where dendrite structures were dissolved, residual laves and δ-phases were found to remain.3. The mechanical properties of the specimens produced in this study is comparable to those reported in the literature for wrought and other AM Inconel 718 specimens.

The mechanical properties of the specimens produced in this study is comparable to those reported in the literature for wrought and other AM Inconel 718 specimens. However, the investigated specimens herein possessed slightly lower yield stress and higher ultimate tensile stress as compared to both wrought and other available data for AM Inconel 718.4. Cracks were found to originate from the surface, and in some cases, near-surface pores in LENS Inconel 718. Cracks grew in a transgranular mode during propagation. Final fracture occurred in a ductile, dimpled mode.5. While fatigue lives are comparable in short-life regime, fatigue lives of LENS Inconel 718 are at least an order of magnitude shorter than its wrought counterpart in the long-life regime.

While fatigue lives are comparable in short-life regime, fatigue lives of LENS Inconel 718 are at least an order of magnitude shorter than its wrought counterpart in the long-life regime. To enhance their fatigue resistance, further improvements need to be made to the additive manufacturing, or post-manufacturing processes, in order to more consistently produce parts with less porosity and more uniform microstructure.AcknowledgmentsResearch was partially sponsored by the Army Research Laboratory and was accomplished under Cooperative Agreement Number W911NF-15-2-0025. The views and conclusions contained in this document are those of the authors and should not be interpreted as representing the official policies, either expressed or implied, of the Army Research Laboratory or the U.S. Government. The U.S. Government is authorized to reproduce and distribute reprints for Government purposes notwithstanding any copyright notation herein.

Government. The U.S. Government is authorized to reproduce and distribute reprints for Government purposes notwithstanding any copyright notation herein. This manuscript was prepared while Nima Shamsaei and Scott Thompson were Assistant Professors at Mississippi State University.References[1]Inconel Alloy 718. Publication no. SMC-045. Special Metals, Sept. 2007. Web. <http://www.specialmetals.com/documents/Inconel%20alloy%20718.pdf>.[2]Chen, Q., N. Kawagoishi, and H. Nisitani. "Evaluation of fatigue crack growth rate and life prediction of Inconel 718 at room and elevated temperatures."Materials Science and Engineering: A 277.1 (2000): 250-257.[3] Ma, Xian-feng, et al. "Fatigue and fracture behavior of nickel-based superalloy Inconel 718 up to the very high cycle regime." Journal of Zhejiang University Science A 11.10 (2010): 727-737.[4] Mercer, C., A. B. O. Soboyejo, and W. O. Soboyejo.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFatigue Behavior of Additive Manufactured 304L Stainless Steel Including Surface Roughness EffectsSeungjong Lee1,2, Jonathan Pegues1,2, Nima Shamsaei1,2*1 Department of Mechanical Engineering, Auburn University, Auburn, AL 36849, USA 2 National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849, USACorresponding Author: shamsaei@auburn.eduAbstractThe fatigue behavior of additive manufactured parts in the as-built surface condition is typically dominated by the surface roughness. However, the fatigue behavior of 304L stainless steel fabricated by laser beam powder bed fusion shows less sensitivity to surface roughness under strain-controlled loading conditions than other additive manufactured materials.

Under force- controlled conditions, however, the high cycle fatigue resistance is much lower for the as-built surface condition than the machined one. This study investigates the underlying mechanisms responsible for fatigue failure for each condition (i.e. strain-controlled or force-controlled). The corresponding cyclic deformation behavior was characterized, and a thorough fractography analysis was performed to identify the features responsible for crack initiation.

The corresponding cyclic deformation behavior was characterized, and a thorough fractography analysis was performed to identify the features responsible for crack initiation. Results indicate that the crack initiation features in both loading conditions are similar, and that the reduced high cycle fatigue resistance for force-controlled fatigue loading compared to strain-controlled one is related to differences in the cyclic deformation behavior of the material.Keywords: Additive manufacturing; Laser beam powder bed fusion (LB-PBF); Surface roughness; Stainless steel; Fatigue;IntroductionMetal additive manufacturing (AM) is a popular technology in high tech industries such as biomedical and aerospace. Even though AM was introduced decades ago, its progress has been lagging due to the lack of advancements in other related technologies such as laser technology and computer aided design (CAD) [1].

Even though AM was introduced decades ago, its progress has been lagging due to the lack of advancements in other related technologies such as laser technology and computer aided design (CAD) [1]. It was only recently that AM has expanded, and metal AM in specific has emerged with several methods: laser beam powder bed fusion (LB-PBF), electron beam powder bed fusion (EB-PBF), and laser beam directed energy deposition (LB-DED) as viable manufacturing techniques [2]. The basic principle of AM is building the component layer- by-layer, thus AM can build complex geometries without the need for assembling multiple components. For the biomedical industry, AM can not only provide implants tailored to an individual patient and a specific injury but also reduce the stiffness of the implants to be more comparable to cortical bones. In the aerospace industry, AM enables light-weighting to improve376fuel-efficiency, using lattice structure or topology optimization.

In the aerospace industry, AM enables light-weighting to improve376fuel-efficiency, using lattice structure or topology optimization. As a result, AM is recognized as an innovative manufacturing technique offering significant advantages over more traditional manufacturing methods.LB-PBF is one of the most common metal AM methods, and its fabricated components have the most competitive mechanical properties among other AM methods [3]. Complex geometries and orientations can be sufficiently supported; hence the chance of collapse is lower than other AM techniques such as DED. In addition, the typical layer thickness of LB-PBF can be as low as 20 (cid:645)m, offering the best surface finish among the various metal AM techniques [4]. However, despite the improved surface finish of LB-PBF materials, the fatigue behavior of as- built parts is still dominated by the surface roughness [5].

However, despite the improved surface finish of LB-PBF materials, the fatigue behavior of as- built parts is still dominated by the surface roughness [5]. Therefore, LB-PBF is studied as a state- of-the-art AM technology especially for fatigue critical applications.Additively manufactured parts have comparable or even better strengths under static loading but often suffer from low or poor ductility. Additionally, under cyclic loading, most AM materials show insufficient performance compared to their wrought counterparts [6], [7], [8]. It has been estimated that more than half of the mechanical failures and fractures are caused by fatigue [9], thus many studies have investigated the fatigue behavior of additive manufactured materials [10], [11]. Yet the underlying effects of the AM process on the resulting structure of the material and their eventual fatigue resistance is not fully established [9], [11], [12].

Yet the underlying effects of the AM process on the resulting structure of the material and their eventual fatigue resistance is not fully established [9], [11], [12]. As such, the adoption of AM across many industries has been limited by their reliability concerns under fatigue loading. Moreover, fatigue behavior is primarily influenced by defects at or near the surface which can initiate and grow cracks leading to premature part failure. It is further impacted by coarse surface roughness that can behave as micro-notches and induce fatigue cracks in the early stages of fatigue life. As a result, the fatigue behavior of AM parts is dependent by surface roughness.304L SS is one of the most common austenitic stainless steels and it has many advantages such as corrosion resistance and a remarkable combination of strength and ductility [9], [12]. Because of these advantages, 304L SS is used in various fields such as nuclear, aeronautical, automotive, and marine industries.

Because of these advantages, 304L SS is used in various fields such as nuclear, aeronautical, automotive, and marine industries. Therefore, LB-PBF AM 304L SS can create a synergy effect from the material itself and its ease of fabrication through AM techniques. However, there is limited information regarding LB-PBF 304L stainless steel (SS), especially in regards to its fatigue behavior. This paper investigates the influence of surface roughness on LB-PBF AM 304L SS under force-controlled and strain-controlled conditions. The fatigue behavior for both as-built and as-machined specimens are presented to understand the effects of surface roughness on the resulting fatigue performance.Experimental ProcedureThe fabrication was conducted using an EOS M290 (EOS, Krailling, Germany) LB-PBF AM machine, and 304L SS powder from LPW Technology Inc. (LPW, Runcorn, UK) with a377reported particle size between 15-45 (cid:645)m.

(LPW, Runcorn, UK) with a377reported particle size between 15-45 (cid:645)m. The EOS M290 was operated with default parameters from EOS itself: laser power of 195 W, the scan speed of 1083 mm/s, the hatching distance of 90 μm, and the layer thickness of 20 μm. Two different geometries were fabricated in the vertical direction including square machining blanks and net-shaped fatigue specimens [13].The surface roughness of the as-built specimen gage sections was investigated using a Keyence VHX-6000 digital microscope (Keyence, Osaka, Japan). The direction of measurement was parallel to both build orientation and load direction. According to the specimen identification number on the top of specimens, the surfaces of the north, west, south, and east sides were measured longitudinally.

According to the specimen identification number on the top of specimens, the surfaces of the north, west, south, and east sides were measured longitudinally. The movement of the recoater arm, in relation to the specimen, is from east to west (from the feedstock bin to the overflow bin) and the direction of gas flow is from north to south (from the back part of the machine to the front) for all specimens. Line surface roughness measurements at 1000x magnification of microscope were taken approximately 2 mm in length along the build direction a minimum of three times and averaged. Figure 1 details the build orientation, the profiled area for line roughness measurement, and the direction of recoater arm and gas flow.NorthNorthWestSouthWest01South~ I !East= 0 ·.c ~ ... = .S:! 0 .... = ='03 line roughnessFigure 1.

= .S:! 0 .... = ='03 line roughnessFigure 1. Schematic build orientation and profiled area for line roughness measurement.Fatigue testing was conducted on an MTS Landmark (MTS, Minnesota, USA) servo- hydraulic test frame with a 100 kN capacity. A set of specimens were tested under force-controlled conditions with the remaining specimens being tested under strain-controlled conditions. For both loading conditions, an extensometer was attached on the side of the gage section to record the constitutive cyclic stress-strain relationship [14].

For both loading conditions, an extensometer was attached on the side of the gage section to record the constitutive cyclic stress-strain relationship [14]. For strain-controlled fatigue testing, three different strain amplitudes of 0.002, 0.003, and 0.0075 (mm/mm) were conducted to incorporate high cycle fatigue (HCF), intermediate cycle fatigue (ICF), and low cycle fatigue (LCF) regimes.378EastFor the force-controlled fatigue testing, the proper loading force was considered from the cross- sectional area and the desired stress amplitude.Fractography for the cross-section of the fracture surface was also conducted using the Keyence VHX-6000. At first, the entire fracture surface was observed with low magnification to estimate overall fracture surface. The higher magnification images were captured to locate districts that contain crack initiation.

The higher magnification images were captured to locate districts that contain crack initiation. Especially, for the as-built specimens, every crack initiated from the surface and therefore, the direction which has the highest surface roughness were more carefully studied.ResultsVarious representative parameters, which describe surface roughness with a numerical value such as (1) the arithmetic mean of all measured values on the profile ((cid:1844)(cid:3028)), (2) the maximum profile height ((cid:1844)(cid:3053)), and (3) the ten-point height of irregularities ((cid:1844)(cid:3053)ISO), were investigated by focal variation digital microscopy. In this study, arithmetical mean deviation ((cid:1844)(cid:3028)) represents the average surface roughness. The average of three measurements from the line-roughness profiles provided the final surface roughness values.

The average of three measurements from the line-roughness profiles provided the final surface roughness values. This was carried out on each north, west, south, and east sides of the gage section with the resulting values shown in Table 1 with their standard deviations in parentheses. Box plots containing Ra values for all directions are shown in Figure 2 showing there is no significant difference in surface roughness for any face. Results show that on average the north face has the highest roughness while the south face has the lowest roughness. However, there is no tendency that the specific side has more surface roughness because each specimen shows different trends across each face.Table 1.

The box graph of surface roughness of each side for each specimenFigure 3 shows the strain amplitude versus reversal to failure (Strain-Life) graph in log- log scale. The results from the as-built condition specimens are shown as red squares and as- machined condition specimens are shown as blue diamonds. The strain-life curve consists of plastic and elastic regions, thus each plastic and elastic curves were calculated first and then combined to get the final strain-life curve. The final strain-life curves are displayed as dotted lines in Figure 3. Table 2 lists the Manson-Coffin coefficients according to the condition of the surface. Manson-Coffin relationship explained as Equation 1 below. The low cycle fatigue (LCF) behavior for both of as-built and as-machined condition were similar, however, the machined condition showed slightly better fatigue resistance. In the high cycle fatigue (HCF) regime, the machined condition showed even better fatigue resistance than the LCF.

Basquin equation(cid:1845)(cid:3015)(cid:3033) = (cid:1827)((cid:1840)(cid:3033))(cid:3003)Where, (cid:1845)(cid:3015)(cid:3033) = fully reversed fatigue strength at (cid:1840)(cid:3033) cycles to failureThe results of fully reversed force-controlled fatigue testing with three different stress levels for both as-built and as-machined specimens are as shown in Figure 4. The two dotted lines represent as the Basquin curves. The coefficients of the blue dotted line are A = 980.8 and B = - 0.1048 and red dotted line are A = 607.2 and B = -3.852E-02. These are calculated by Matlab using Basquin equation as shown in Equation 2. Figure 4 reveals that, under force controlled conditions, the material is much more sensitive to surface roughness than when under strain- controlled conditions.

Figure 4 reveals that, under force controlled conditions, the material is much more sensitive to surface roughness than when under strain- controlled conditions. For example, when the stress amplitude was 350MPa for the as-machined specimen, 2(cid:1840)(cid:3033) was already over 2,000,000 reversals while the as-built condition only reached 10,000 reversals, a difference of 2 orders of magnitude. Therefore, the range of stress amplitude for machined specimens was 350 MPa ~ 400 MPa and as-built specimens was 200 MPa ~ 375381MPa. In addition, there is a compelling distinction between the two surface conditions such as intersection (A) between two trend lines and slope (B).l03 ,----~~--~~----.-,",---~===== I □ As-built I ◊ As-machined !Force-controlled fatigue testing....□□l02 ~ - - - -~ - - - -~ - - - -~ - - - -~ - - - -~ l08l03104l05l06107Reversal to Failure, 2NfFigure 4.

Stress amplitude (MPa) versus reversal to failure (2(cid:1840)(cid:3033)) graph of force-controlled fatigue testing.Figure 5. Fracture surfaces of strain-controlled tested as-built condition specimen #04. The strain amplitude was 0.002 (mm/mm) and the reversals to failure was 139,312.382The fractography analysis provides details regarding the fracture surface such as crack initiation and crack propagation. The observed fractographic image of specimen # 04, shown in Figure 5, details many of the common features observed for as-built specimens. This specimen was tested in the HCF regime at a strain amplitude of 0.002 (mm/mm). The fracture surface indicates that multiple cracks initiated on the north surface, coalesced into a single dominant crack, and grew through ~98.0% of the cross-sectional area before final rupture.

The fracture surface indicates that multiple cracks initiated on the north surface, coalesced into a single dominant crack, and grew through ~98.0% of the cross-sectional area before final rupture. Figure 6 shows cross- sectional fracture surfaces two specimens tested under strain-control at (cid:632)a = 0.003 (mm/mm) for (a) as-built condition specimen # 37 which has 47156 reversals to failure and (b) as-machined condition specimen # 18 which has 64080 reversals to failure. For the as-built condition, multiple cracks initiated from the southern side of specimen. On the other hand, the as-machined specimen had the dominant crack initiate on the surface, however, several cracks also initiated sub-surface.Figure 7 shows the fracture surfaces of two specimens tested under force-control at 375 MPa in the (a) as-built condition specimen # 13 and (b) as-machined condition specimen # 30.

The crack initiation under force-controlled conditions show similar characteristics as strain-controlled fatigue such as multiple cracks initiating on the surface of the as-built condition and sub-surface cracks initiating on the as-machined condition. In addition, there is a striking difference in the crack growth behavior between strain-controlled fatigue as shown in Figure 6 and force-controlled fatigue as shown in Figure 7. While the strain-controlled tests showed stable crack growth through the majority of the specimen cross-section, the force-controlled tests fail abruptly at approximately 40% of the cross-sectional area.Figure 6. (a) Fracture surfaces of strain-controlled tested as-built condition specimen # 01. (b) Fracture surfaces of strain-controlled as-machined condition specimen # 18. For both specimens,383entire surface images were taken by 300X magnification and zoomed images were taken by 1000X magnification.

For both specimens,383entire surface images were taken by 300X magnification and zoomed images were taken by 1000X magnification. Strain amplitude was 0.003 (mm/mm) for both of them.Figure 7. (a) Fracture surfaces of force-controlled tested as-built condition specimen # 13. (b) Fracture surfaces of force-controlled as-machined condition specimen # 30. For both specimens, entire surface images were taken by 300X magnification and zoomed images were taken by 1000X magnification. Stress amplitude was 375 MPa for both.DiscussionIn this study, the tests in the as-built condition always showed cracks initiating from the rough surface, however, the actual crack initiation regions were often different from the highest surface roughness region. All of the crack initiation regions were investigated by fractography and compared with surface roughness data (Table 1).

All of the crack initiation regions were investigated by fractography and compared with surface roughness data (Table 1). Arithmetical mean deviation ((cid:1844)(cid:3028)) can explain the surface roughness but it does not give the critical information needed to determine the crack initiation point. Maximum profile valley ((cid:1844)(cid:3049)) was also investigated as a different approach, but the results were also not sufficient to identify the critical location of the specimens. This is mostly due to the roughness values not capturing the geometry of the micro-notches. Previous studies on surface roughness effects of additive manufactured Ti-6Al-4V has shown that sufficient predictions can be made by incorporating a geometric term into the effective stress concentration factor estimation [15].There are many factors that affect surface roughness such as the direction of a recoater arm, distribution of powder particle’s size, and gas flow [16], [17], [18].

First, the powder was diffused by the recoater arm from the feedstock bin (east) to the overflow bin (west) between layers, hence384the height of each layer can be different because the blade on the recoater arm extrude the powder. Second, the variations in particle sizes can result in changes in the surface condition. If there is a large particle at the end of the part, this particle will remain as an un-melted powder on the surface because of the insufficient energy produced by laser process. This discrepancy is directly related to the distribution of powder. Third, the inert gas such as argon or nitrogen flow in the machine from north to south to decrease oxygen content during the printing. Some powder is ejected into the air when the laser interacts with the powder bed, at the same time the gas flow can carry these scattered powders which can disturb the powder bed near an unfused part or adhere them to a previously fused part.

In the latter case the subsequent spreading of the powder can lead to an uneven layer of powder near the adhered particle and result in poor fusion at the surface or even within the part. In addition, the gas flow can become turbulent, especially at the corners of the plate resulting in the plume not being carried directly away from the laser beam path. When the plume interacts with the laser it can slightly defocus the laser leading to larger/smaller spot sizes and insufficient melting. This complicated gas flow gives rise to a higher probability of surface roughness variation as well as internal defects.The main difference between fracture surfaces of as-built and as-machined specimens is the region of crack initiations. For as-built specimens under both strain-controlled and force- controlled fatigue, cracks are initiated from multiple places on the surface. They grow separately at the beginning, then coalesce leading to a single dominant crack.

They grow separately at the beginning, then coalesce leading to a single dominant crack. On the other hand, for as- machined specimens under both conditions, the main crack initiation point is at the surface and does not appear to be influenced by other cracks at the surface but does coalesce with cracks initiated from internal defects, as shown in Figure 6 (b) and Figure 7 (b).The specimens tested in force-control were revealed to be more sensitive to surface roughness than specimens tested in strain-control. For strain-controlled fatigue tests as shown in Figure 6, the crack growth is mostly similar between as-built and as-machined surface conditions. The cracks propagated up to 97.50% for (a) and 97.01% for (b). Under strain-controlled conditions the force to maintain the actual strain of the specimen reduces as the strain accumulates allowing the crack to remain stable throughout the entire fatigue life.

Under strain-controlled conditions the force to maintain the actual strain of the specimen reduces as the strain accumulates allowing the crack to remain stable throughout the entire fatigue life. On the other hand, for force-controlled fatigue tests as shown in Figure 7, the crack propagated to 30.36% for (a) and 45.30% for (b). The force-controlled fatigue testing maintains the force, thus increasing the actual stress as the crack develops and eventually leads to unbounded deformation. Thus, in the as-built condition this occurs suddenly as the multiple crack interact and coalesce leading to unstable crack growth.

Thus, in the as-built condition this occurs suddenly as the multiple crack interact and coalesce leading to unstable crack growth. Conversely, under stain-control conditions the deformation in each cycle is limited by the loading condition and thus the crack grows at a stable rate regardless of its interaction with other cracks.ConclusionsThe effects of surface roughness of the as-built condition of LB-PBF 304L SS AM under strain-controlled and force-controlled fatigue testing were investigated in this study. For385comparison, as-machined specimens were also fabricated and tested under the same conditions. As a result, four different fatigue life curves and the fractography images for each condition were obtained. The following conclusions can be made as a result of the findings:1. Surface roughness can vary with no obvious correlation to direction.

The following conclusions can be made as a result of the findings:1. Surface roughness can vary with no obvious correlation to direction. Even though the specimen has cylindrical design, surface roughness of each sides (north, west, south, and east) show variations.2. The mechanical debit on fatigue life as-built and as-machined specimen under strain-control is much smaller than under force-control.3. Multiple cracks initiated on the surface of the as-built specimens. However, there is onlyone crack initiation close to the surface of as-machined specimens.4. Crack growth stability of force-controlled conditions is much lower for a propagating crack than strain-controlled conditions resulting in the observed discrepancy in fatigue resistance for the loading conditions.AcknowledgementThis material is based upon work supported by the National Science Foundation (NSF)under Grant No. 1657195.References[1][2]T. Wohlers and T.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFatigue behavior of additive manufactured parts in different process chains – An experimental studyEckart Uhlmanna, Georg Gerlitzkya, Claudia FleckbaInstitute for Machine Tools and Factory Management (IWF), TU Berlin, Pascalstraße 8-9, Berlin, Germany bDepartment of Materials Engineering, TU Berlin, Straße des 17. Juni 135, Berlin, GermanyAbstractMetal based Additive Manufacturing (AM) has experienced dynamic growth in recent years. However, the global distribution of Additive Manufacturing is limited by the fact the produced parts suffer from bad surface quality and the material properties concerning fatigue life are still an object of current investigations which limits possible applications of AM parts.

Due to this fact metal AM processes are often followed by a post process to ensure a better surface quality. In this paper the authors present results where fatigue life and different post processes of additive manufactured parts are investigated. Subsequently, surface roughness, high cycle fatigue, fracture behavior and microstructure have been characterized. Finally the results for the different post processing states have been compared and surface properties as well as microstructure have been correlated with the fatigue properties in order to evaluate how different process chains influence the High cycle fatigue (HCF) behavior of additive manufactured parts.IntroductionFreeform fabrication processes such as Selective Laser Melting (SLM) offer high geometrical freedom compared to conventional manufacturing processes like turning milling or grinding.

In addition the relevance of Selective Laser Melting is steadily growing as it applicable to a wide range of materials and it empowers the fast production of prototypes and serial components. Therefore, Selective Laser Melting delivers an answer to the continuously increasing complexity of serial components and enables the resource saving manufacturing of complex parts and hard to machine components [1]. Especially in the aviation industry SLM is already successfully adapted for serial components due to the expensive materials used and the high material removal of conventional manufacturing processes. At the moment however SLM is only cost efficient in certain business cases and is preferred at low production volumes, high raw material costs and at the processing of hard to machine materials.

At the moment however SLM is only cost efficient in certain business cases and is preferred at low production volumes, high raw material costs and at the processing of hard to machine materials. This is mainly due to the fact that the fixed costs of the SLM process are lower than the fixed costs of comparable conventional manufacturing processes while the recurring costs of the SLM processes are higher than conventional manufacturing processes [2]. Currently the main problems concerning Additive Manufacturing processes are the achievable surface quality, geometrical accuracy, uncertainty about mechanical properties and high costs [3]. Due to these problems SLM manufactured parts often need to be post processed with manufacturing processes such as turning, milling, blasting or vibratory finishing. In this paper the authors present a work which unites different post processes and fatigue life of additive manufactured parts.

In this paper the authors present a work which unites different post processes and fatigue life of additive manufactured parts. First, parts are produced with Selective Laser Melting with constant process parameters such as laser power or scanning speed. After this the parts are post processed with48vibratory finishing and turning followed by an investigation concerning roughness, high cycle fatigue (HCF) and fracture behavior. Finally the results for the different post processing states have been compared and surface properties and microstructure have been correlated with the fatigue properties in order to evaluate how different process chains influence the HCF behavior of additive manufactured parts.Experimental SetupAll parts were produced on the SLM machine SLM 250 HL, MTT Technologies GmbH, Lübeck, located at the Production Technology Center in Berlin, Germany.

To process the specimens stainless steel powder 316L (1.4404) was used with a typical grain size distribution of 20 to 63µm. The powder properties were obtained from the suppliers’ certificate following European standard DIN EN 10204. The chemical composition of the powder is given in Table 1. A scanning electron microscopy (SEM) image of the powder is shown in Figure 1.Table 1:Chemical composition of processed powderElement Mass percentageFe Bal.C 0.02Si 0.8Mn 1.7Cr 17.7Mo 2.2Ni 10.5The image illustrates that not all powder particles have a spherical form which might result in a lower flowabilty of the powder. Due to the fact that stainless steel powder normally offers high flowabilty rates the shown powder sizes should also offer an acceptable flowability rate.

Due to the fact that stainless steel powder normally offers high flowabilty rates the shown powder sizes should also offer an acceptable flowability rate. This assumption was validated with Hall-flowability measurements which showed an average flow time of 3.4 s.49Figure 1: SEM image of processed stainless steel powderThe parts were produced with optimized building parameters regarding density of the part and process time, provided by the machine supplier and are shown in table 2. These parameters are commonly used for the SLM process and are therefore used as reference parameters for the different process chains. All specimens were oriented vertically on the building platform.Table 2:Overview of SLM process parametersProcess parameters Layer thickness Laser power Scanning speed Hatch pattern Hatch distance0.05mm 275 W 760 mm/s Chess 0.12mmTest specimen Two different specimen geometries were produced for the different post processes.

Specimen A was used for the turning process. It is clearly visible that this specimen significantly differs from specimen B which was used for the vibratory finishing process and the SLM shape HCF50investigations. This is mainly due to the fact that the part has to be longer in order to be clamped inside the machine tool. The dimensions are illustrated in Figure 3.Specimen A was produced with SLM and post processed with turning which lead to specimen D (figure 2). The building direction is indicated with a black arrow. Specimen B was produced with SLM with no further post process. Specimen C was produced with SLM followed by an additional vibratory finishing process.Figure 2: Dimensions and illustration of different specimensThe post processes were selected because they are commonly used as post processes for SLM parts in industry and research.

Especially vibratory finishing could be a promising finishing process for SLM parts because the accessibility of geometrical complex structures of SLM parts is possible due to the small abrasive media used. The turning process was carried out on a CTX gamma 1250 CT machine from DMG Mori Europe Holding AG, Winterthur, Switzerland. The vibratory finishing process was carried out as a two stage process on a MF 8/1, Multi finish GmbH & Co. KG, Schömberg. In a first process step triangular abrasive media was used for a process time of 120 minutes. In a second process step cylindrical abrasive media was used for a process time of 60 minutes. The process parameters for the turning and the vibratory finishing process are given in table 3 and table 4.Table 3:Process parameters of the turning processTurning process Depth of cut ae Feed f Cutting speed vc0.5 mm 0.15 mm/ rev.

140 m/minTable 4:Process parameters of the vibratory finishing processVibratory finishing process Processing time Abrasive media Excitation frequency fe1st process step 120 min Triangular 50 Hz2nd process step 60 min Cylindrical 50 Hz51Microstructural investigations For the microstructural characterization, longitudinal and cross-sections for each kind of specimen were prepared. The longitudinal sections contain the area of the fracture origin and the final fracture; the cross-sections are situated about 1 cm below the fracture surface. After grinding and polishing, the sections were etched with V2A-etchant to develop the microstructureRoughness measurements All surface roughness measurements were conducted with the tactile measurement system Hommel Nanoscan etamic 855, Jenoptik, Jena. The investigated target values were Ra and Rz. All specimens were measured three times and turned by 120 degrees to ensure reproducibility for all measurements.

The investigated target values were Ra and Rz. All specimens were measured three times and turned by 120 degrees to ensure reproducibility for all measurements. The average value of these three measurements was taken to obtain the target value of one specimen. In addition, the mean standard deviation was calculated for all measurements.High cycle fatigue (HCF) test series High cycle fatigue tests were performed with a rotating bending machine (Carl Schenck AG, Darmstadt, type Rapid Punz) at a frequency of 100 Hz under air cooling to a maximum number of cycles (Nmax) of 107. Two surface stress amplitudes were chosen in the HCF range for each finishing state to yield numbers of cycles to failure (Nf) of about 5x105 and 2x106.

Two surface stress amplitudes were chosen in the HCF range for each finishing state to yield numbers of cycles to failure (Nf) of about 5x105 and 2x106. Further specimens were tested at lower surface stress amplitudes, in the transition range to the endurance limit.SEM analysis The fracture surfaces of selected specimens, all tested with surface stress amplitudes leading to failure within 106 cycles, were investigated by scanning electron microscopy (SEM; CamScan REM Serie 2, Obducat, Lund, Sweden) at an accelerating voltage of 20 kV in the secondary electron mode. The specimens were cut about 1 cm below one of the fracture surfaces under continuous water-cooling while the fracture surfaces were protected by a polymer cover, and imaging was performed after ultrasonic cleaning in ethanol.Comparison with state of the art fatigue testing Fatigue testing of additive manufactured products has already been investigated by some researchers e.g. [4], [5], [6].

[4], [5], [6]. The main testing method was the four point bending test as in this work. Often the main focus was to compare different heat treatment strategies or build orientations to ensure the best possible fatigue life [5], [6]. In some works even a concrete product was produced with AM and successfully dimensioned in terms of fatigue life [4]. In other work the focus was to develop the best possible process parameters for the highest possible fatigue life [7]. The results in this work correlate well with other works regarding the dependency of roughness and fatigue life. A unique feature of this work is to reveal the differences between the examined process chains regarding fatigue life as well as different fracture behaviors. This should be helpful to identify reasonable holistic process chains for AM. Initially 60 specimens were produced with SLM followed by the different post processes turning and vibratory finishing.

This should be helpful to identify reasonable holistic process chains for AM. Initially 60 specimens were produced with SLM followed by the different post processes turning and vibratory finishing. In total 20 of each kind of specimens (B, C and D) were produced. Subsequently, for all specimens the surface roughness was determined. The microstructure and the fracture behavior of one typical specimen of each kind was investigated by light and scanning52electron microscopy (LM, SEM), respectively. After this, all specimens were tested regarding their high cycle fatigue behavior on the Punz machine.Results/ DiscussionMicrostructural investigationsFigure 3 shows the typical microstructure of the SLM manufactured specimens. The building direction is indicated with an arrow, in this case perpendicular to the image plane. It consists of an austenitic microstructure. The cross-sections also nicely represent the hatch pattern of the manufacturing process.

It consists of an austenitic microstructure. The cross-sections also nicely represent the hatch pattern of the manufacturing process. Each of the longitudinal structures is made up of many grains which lack the polygonal structure usually observed in austenitic steels. In some areas, dendrites are visible. All specimens contain binding defects, visible as black voids.abFigure 3: (a) Austenitic microstructure; (b) Scanning pattern of a turned specimenThe surface of the turned specimen is smooth with some voids present. The surface of the vibratory finished specimen is also very smooth, and the microstructure just below the surface looks plastically deformed. Nevertheless, more shallow defects are visible on the surface than on the turned surface. The SLM shape specimen has a very rough surface.

Nevertheless, more shallow defects are visible on the surface than on the turned surface. The SLM shape specimen has a very rough surface. Numerous large defects and molten and re-solidified powder particles are visible and indicated with white arrows in figure 4(b).53cbaFigure 4: Comparison of differently post processed surfaces (a) Vibratory finished (b) SLM-nearnet shape (c) TurnedThe longitudinal section in figure 5 a2 shows a secondary crack that runs through a binding defect just below the surface. We may assume that this fatigue crack originated at this defect but could not develop into a dominant crack due to higher crack growth in the level of the later fracture surface. Nevertheless, more shallow defects are visible on the surface than on the turned surface.54Figure 5: LM: cross-sections (left row, “1”) and longitudinal sections (right row, “2”) of a) turned, b) vibratory finishing, c) SLM shape specimens.

Note the different scale of b2 as compared to a2 and c2.Roughness measurements1640Arithmeticalmeandeviationofthe assesedprofileRa8µmSlmshapeSpecimenVibratoryfinishingTurned13,2851,7371,076Tool:SelectiveLaser Melting-MachineMTT -SLM 250HLMaterial:Stainlesssteel316LFigure 6: Arithmetical mean deviation of the assessed profile Ra of specimens55100250Maximum heightofprofileRz50µm82,86317,0334,957SlmshapeSpecimenVibratoryfinishingTool:SelectiveLaser Melting-MachineMTT -SLM 250HLMaterial:Stainlesssteel316LTurnedFigure 7: Maximum height of profile Rz of specimensFigure 6 and Figure 7 show the measured target values Ra and Rz for the different specimens. The SLM shape specimens have by far the highest surface roughness of all specimens with a roughness Ra of 13.29 and Rz = 82.86. The vibratory finished specimens have a mean arithmetic surface roughness of Ra = 1.74 and Rz = 17.03. Turned specimens showed the best surface roughness with Ra of 1.08 and Rz 4.96.

The vibratory finished specimens have a mean arithmetic surface roughness of Ra = 1.74 and Rz = 17.03. Turned specimens showed the best surface roughness with Ra of 1.08 and Rz 4.96. The vibratory finished specimens have a significantly higher surface roughness than the turned specimen for both target values Ra and Rz. Better surface qualities for the vibratory finishing process could be obtained with a longer process time and special abrasive media, due to the fact that fatigue resistance is highly dependent on the surface roughness [8]. This may lead to a low fatigue resistance of the SLM shaped specimen, with the turned specimen possibly showing the best fatigue resistance. The measurements also revealed the highest standard deviation for SLM shape specimen. The high surface roughness is mainly due to the fact that the typical particle size distribution for the SLM process is 20 to 63 µm. This limits the achievable surface quality.

The high surface roughness is mainly due to the fact that the typical particle size distribution for the SLM process is 20 to 63 µm. This limits the achievable surface quality. With a smaller particle size distribution better surface qualities would be achievable but this contains other risks and problems such as respirability of the finer powder and worse flowability of the powder. The higher standard deviation has two possible reasons. First, the poor reproducibility of SLM manufactured parts, because of different temperatures in the melt pool which are caused by a non-symmetrical protective gas flow and the location of the part on the building platform during the process which causes different heat transfer coefficients. The second reason is the measurement systems. Due to the high surface roughness optical as well as tactile measurement systems reach their limits [9].

The second reason is the measurement systems. Due to the high surface roughness optical as well as tactile measurement systems reach their limits [9]. Optical measurement systems struggle with reflecting surfaces while the stylus tip of a tactile measurement can break due to possible undercuts and pores of the selective laser melted part.High cycle fatigue (HCF) test series Figure 8 illustrates typical striation due to the cyclic loading while Figure 9 shows the results of the rotating bending tests. The dashed lines in the HCF regime are a guide for the eye to highlight the fatigue life ranges of the differently finished SLM specimens. The gray arrows denote specimens that did not fail at cycle numbers above the defined maximum number of cycles to failure (Nmax, denoted by the bold black line). The SLM shape specimens exhibit the lowest fatigue life resistance.56Figure 8: Striation of HCF specimensAfter vibratory finishing or turning, the life expectancy increases clearly.

The SLM shape specimens exhibit the lowest fatigue life resistance.56Figure 8: Striation of HCF specimensAfter vibratory finishing or turning, the life expectancy increases clearly. The differences seem to be slightly higher for higher surface stress amplitudes, and slightly lower for stress amplitudes in the transition range. The SLM shape and vibratory finished specimens show a huge scatter in the fatigue life, specifically in the transition range, with specimens failing at around Nf = 2x106 while others survive Nf = 2x107 at the same surface stress amplitude. For the turned specimen, the numbers of cycles to failure scatter much less in this fatigue range. The results of the fatigue tests correlate well with the differences in surface roughness between the specimen types. Local height variations on the surface originating from the powder particle size, as in the SLM shape specimens, or from turning marks as in the turned specimens, act as stress raisers.

Local height variations on the surface originating from the powder particle size, as in the SLM shape specimens, or from turning marks as in the turned specimens, act as stress raisers. Fatigue failures usually originate from the surface; furthermore, in rotating bending tests the highest tensile stresses act at the surface. Rougher surface states result in locally even higher stresses, corresponding with lower numbers of cycles to failure. The higher fatigue endurance of the vibratory finished and turned specimens may further be due to compressive residual stresses originating from the finishing procedures. Further, defects just below the surface might be compressed by plastic deformation near the surface. Especially in [5] the same material was tested with the same testing method and the same orientation on the building platform. The results correlate as the specimen failed between 106 and 107cycles and a maximum stress at knee point of 180 MPa for as built specimens.

The results correlate as the specimen failed between 106 and 107cycles and a maximum stress at knee point of 180 MPa for as built specimens. In that paper also a significant increase through machining the specimens was observed. However results may differ due to the fact that a different SLM machine and different machining techniques were used. In addition it is hard to compare different investigations regarding fatigue because different materials, surface integrity and SLM machines are investigated.57105106107= Nmax10850100150200250300Surface stress amplitudesa(MPa)number of cycles to failure NfTurnedVibratoryfinishedSLM shapeFigure 9: S, N-curves for SLM-316L specimensSEM analysis Typical fracture surfaces of specimens that failed at about 106 cycles are shown in Figure 10. The surveys in the upper row show the smoother fatigue and rough final fracture areas.

The surveys in the upper row show the smoother fatigue and rough final fracture areas. All specimens exhibit several failure origins, denoted by arrows: The continuous arrows with a bold head point to the one (a1, b1) or two (c1) main, dominant fracture origin(s), and the dashed arrows with a line-head point to additional non-dominant origins. The higher magnification micrographs in the bottom row show the area of the dominant, or of one of the dominant fatigue origins on the fracture surfaces together with the specimen surfaces nearby. All fatigue failure(s) originated at binding defects at (a2, b2) or just below (c2) the surface.

All fatigue failure(s) originated at binding defects at (a2, b2) or just below (c2) the surface. Please note that the magnified sections of the vibratory finished and the turned specimen are oriented in the same direction as the survey above while the SLM shape specimen has been rotated clock-wise by 90 to acquire the magnified view.58Figure 10:SEM micrographs of fracture surfaces of specimens with different surfacefinishing: a) turned, b) vibratory finishing, c) SLM shape; all loaded with surface stress amplitudes leading to numbers of cycles to failure in the range of Nf = 106While the SLM shape specimens exhibit a rough surface, with rounded powder particles clearly visible (a2; the white dashed line serves as a guide to the eye to discriminate the fracture and specimen surfaces above and below the line), the vibratory finished specimens have a relatively even and smooth surface (b2). Turning marks are visible on the surfaces of the turned specimens (c2).

Turning marks are visible on the surfaces of the turned specimens (c2). Inner defects just below the surface act as local stress raisers; together with the different surface roughness features and possible compressive residual stresses in the surface finished states, they define the local stress state, and failure starts where the local tensile stresses are highest (comp. figure 9a2).The final fracture surfaces show dimples, as to be expected for the inherently ductile austenitic materials, and occasional fatigue striations were observed on the fatigue fracture surfaces (data not shown). The black areas on the micrographs are due to charging effects of only partially attached or less electron-conductive particles on the surfaces and fracture surfaces, usually found within depressions.ConclusionIn this work the influence of two different post processes for Selective Laser Melted specimens regarding their fatigue behavior was investigated.

The investigated material was stainless steel 316L. The approach showed that the bad surface quality of additive manufactured parts is a key problem of this technology, not only for functional surfaces, but as well for the fatigue life. Turned specimens showed the best fatigue behavior while SLM manufactured specimens showed the worst. This correlates well with the obtained surface roughness, also the data for fatigue life of SLM manufactured and vibratory finished specimens scatter a lot. However the results also showed that even though overall a high material density exists, small59defects such as unmolten powder particles lead to dominant cracks. These results demonstrate the importance of integrating the SLM process into an appropriate and cost efficient process chain to increase the fatigue life of additive manufactured parts and spread the applications possibilities of SLM.

In the future hard to machine materials, produced by SLM, such as titanium or nickel based alloys which are used for dynamically loaded operations in aviation should be investigated regarding different process chains. In addition the conventional process chain should be compared with the additive manufacturing process chain regarding fatigue life of the produced parts. Concluding, other process chains such as blasting should be investigated and these results should be combined with existing results regarding heat treatment methods.References [1] Meiners, W. Direktes Selektives Laser Sintern einkomponentiger metallischerWerkstoffe, Dissertation, Shaker Verlag, Aachen, 1999.[2] W. Frazier. Metal Additive Manufacturing: A review. Journal of Materials Engineering[3][4][5][6][7][8]and performance, Vol 23 ,6, 1917-1928, 2014. E. Yasa, J-P. Kruth, J. Deckers. Manufacturing by combining Selective Laser Melting and Selective Laser Erosion/laser re-melting.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFatigue Behavior of Laser Beam Directed Energy Deposited Inconel 718 at Elevated TemperatureAlexander S. Johnson1, Rakish Shrestha2,3, P.D. Nezhadfar2,3, Nima Shamsaei2,3*1TriVector Services Incorporation, Huntsville, AL 2Department of Mechanical Engineering, Auburn University, Auburn, AL 3National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, ALCorresponding author: shamsaei@auburn.eduAbstractNickle based super alloys such as Inconel 718 are being extensively used to manufacture turbine blades for jet engines due to their superior mechanical properties at higher working temperatures.

Furthermore, poor machinability associated with Inconel 718 also makes it an attractive material for additive manufacturing processes, which possess the capability to fabricate near net shaped parts. Hence, in this study, the fatigue behavior of Inconel 718 fabricated using laser beam directed energy deposition (LB-DED) is investigated under strain-controlled, fully- reversed conditions at an elevated temperature of 650° C. Fractography analysis was conducted to determine the failure mechanism for additive manufactured Inconel 718 due to higher working temperatures. The results obtained from the fatigue and fractography analysis were then compared with the results obtained from fatigue tests conducted at room temperature. At elevated test temperature, LB-DED Inconel 718 specimens exhibited lower fatigue resistance compared to the tests conducted at the room temperature, primarily in the low cycle fatigue regime.

At elevated test temperature, LB-DED Inconel 718 specimens exhibited lower fatigue resistance compared to the tests conducted at the room temperature, primarily in the low cycle fatigue regime. Whereas, in the high cycle fatigue regime the effect of test temperature was observed to be minimal. Furthermore, secondary cracks resulting from the formation of brittle behaving precipitates on the grain boundaries was also evident from the fractography analysis indicating significant changes in the microstructural features of LB-DED Inconel 718 as a consequence of elevated test temperature.Keywords: Inconel 718, Fatigue, Elevated temperature, Laser beam directed energy depositionIntroductionInconel 718, a nickel based super alloy, with superior mechanical properties and corrosion resistance in a wide range of temperatures (i.e.

from ambient temperature to elevated temperatures) are being extensively used in nuclear reactors and in the fabrication of turbine blades for jet engines [1, 2]. Along with its superior mechanical properties, Inconel 718 also possess high toughness, which makes this alloy difficult to be machined [3, 4]. As a result, fabrication of these materials using additive manufacturing (AM) techniques is recently gaining attention due to the ability of AM process to manufacture near-net shaped parts, which would eliminate/decrease the need for machining.Despite various advantages of AM processes, parts fabricated using these techniques also come with some drawbacks, which includes, presence of defects resulting from lack of fusion (LoF) between the subsequent layers and gas entrapped pores [5]. Furthermore, due to high cooling584rates inherent to AM processes, the microstructural features also vary greatly when compared with those of wrought counterparts [6].

Furthermore, due to high cooling584rates inherent to AM processes, the microstructural features also vary greatly when compared with those of wrought counterparts [6]. Therefore, to ensure the reliability of additive manufactured (AM) Inconel 718 parts, effect of defects and microstructural features on the fatigue behavior of this alloy at room and elevated temperatures need to be thoroughly understood.Johnson et al. [7] studied the uniaxial strain-controlled, fully reversed room temperature fatigue behavior of Inconel 718 fabricated using laser beam directed energy deposition (LB-DED) process. Results showed that the employed heat treatment procedure was capable of transferring large elongated grains, resulting from directional heat dissipation during manufacturing process, into finer and more equiaxed grain structure.

However, LB-DED Inconel 718 exhibited a lower fatigue resistance as compared to the wrought counterpart, which was attributed to the presence of brittle inclusions of metal-carbide/oxide and gas entrapped pores located near the surface of the specimens.Konecna et al. [8], studied the effect of layer orientation on the resulting room temperature fatigue behavior of near-net shape (i.e. as-built surface condition) Inconel 718 manufactured using laser beam powder bed fusion (LB-PBF) process. This study showed that the crack always initiated from pore-like defects close/on the surface of the specimens. Furthermore, a good correlation between the surface roughness values and the resulting fatigue lives was reported as specimen with higher surface roughness always exhibited shorter fatigue lives.

Furthermore, a good correlation between the surface roughness values and the resulting fatigue lives was reported as specimen with higher surface roughness always exhibited shorter fatigue lives. Some anisotropy in the fatigue behavior of LB-PBF Inconel 718 was also evident as the lowest fatigue resistance was found to be obtained for specimens with loading direction parallel to the build direction.Yadollahi and Shamsaei [6] reported a minimal effect of surface roughness on the fatigue behavior of LB-PBF Inconel 718 subjected to uniaxial force-controlled loading in laboratory environment (i.e. 26°C). Similarity in fatigue resistance of machined and as-built (non-machined) specimens was associated to the migration of large sub-surface defects that were brought to the surface through the machining process. Such defects acted similar to the surface roughness in the as-built specimens resulting in the similar fatigue lives.

Such defects acted similar to the surface roughness in the as-built specimens resulting in the similar fatigue lives. Hence, this study highlighted the importance of depth of machining and its effect on the resulting fatigue behavior of AM parts.Recently, many studies have focused on investigating the influence of defects and microstructure on the room temperature fatigue behavior of AM Inconel 718. However, to the best of authors’ knowledge there is no study on the effect of defects and microstructural features on the fatigue strength of AM Inconel 718 at elevated temperatures in the literature. Therefore, in this study, the fatigue behavior of LB-DED Inconel 718 is investigated under uniaxial strain-controlled fatigue test at elevated temperature of 650°C. The experimental procedure for fabricating parts and conducting fatigue tests are explained. Results obtained from the experimental tests are presented.

The experimental procedure for fabricating parts and conducting fatigue tests are explained. Results obtained from the experimental tests are presented. Eventually, based on the results, important conclusions are drawn.Experimental ProcedurePlasma rotate electrode processed (PREP) Inconel 718 powder with size distribution ranging between 45 to 150 μm were used to fabricate parts in an OPTOMEC LENSTM 750, an LB- DED system. Prior to the fabrication of specimens, preliminary line test was conducted to determine appropriate process parameters capable of producing desired geometry as well as good585fusion with the build plate [7]. Based on the results obtained from the line test, laser power of 350 W, scan speed of 15 mm/s, powder flow rate of 0.057 g/s, and hatch spacing of 0.529 mm were proposed as the optimal process parameters [7]. Cylindrical rods with diameter of 8 mm and 80 mm length were fabricated one at a time in vertical direction (i.e.

Cylindrical rods with diameter of 8 mm and 80 mm length were fabricated one at a time in vertical direction (i.e. perpendicular to the build plate) in an argon purged environment using the proposed process parameters. Cylindrical rods were fabricated instead of net shaped parts in order to ensure that sufficient material is removed during the machining process and minimize the effect of surface roughness. This is important as previous studies have shown that when insufficient material is removed, machined specimens exhibits similar fatigue resistance to that compared to the as built specimen [6].Cylindrical rods were further machined into specimens with uniform gage section following ASTM Standard E606 [9] with the final dimensions shown in Fig. 1. It also needs to be mentioned that the gage diameter of the specimen was slightly deviated from the minimum recommended dimension of 6.35 mm by ASTM Standard E606 [9] to minimize the machining process.

Machined samples were further subjected to standard heat treatment procedure for Inconel 718 reported in [10]. Specimens were first heated up to 940°C for 2 hours followed by air cooling at the rate of 50°C per hour. Once the specimen reached to 621°C, the temperature was then held constant and the specimen was aged for 8 hours followed by air cooling.~ ,a 1 t--rl...____ _ 15mm1 --r i - ---1 26.85mm ~ -I_____.+ t2-3.50mm I I -@) -1) 7mmt-80mmFigure 1: Specimen geometry of LB-DED Inconel 718 alloy used for fatigue testing.Strain-controlled fatigue tests were conducted in an MTS Landmark servo-hydraulic test frame with load cell of 100 kN. Furthermore, the test frequency applied was calculated based on the strain amplitude to obtain similar average strain rate for all of the fatigue tests. MTS high temperature extensometer was used to monitor the deformation on the gage section of the specimen as well as to control the fatigue tests.

MTS high temperature extensometer was used to monitor the deformation on the gage section of the specimen as well as to control the fatigue tests. All of the mechanical tests at the elevated temperature were conducted in an MTS high temperature chamber at 650°C using a sinusoidal wave form. In order to only heat the gage section of the specimens, grips were cooled using running water during the fatigue test as shown in Fig. 2, which illustrates the complete test setup. Fractography analysis was conducted using scanning electron microscope (SEM) to investigate factors responsible for crack initiation and the failure mechanism at elevated temperatures.586Load Cell -----High Temperature Chamber ------- High Temperature Extensometer-- Water-cooled GripsFigure 2. High temperature fatigue test setup showing all of equipment used in the study.Experimental Results and DiscussionsUniaxial strain-controlled fatigue test results presented in Fig.

High temperature fatigue test setup showing all of equipment used in the study.Experimental Results and DiscussionsUniaxial strain-controlled fatigue test results presented in Fig. 3 show some effect of testing temperature (i.e. room temperature vs. high temperature) on the fatigue behavior of LB-DED Inconel 718, which primarily depended on the fatigue life regimes. As it can be clearly seen from Fig. 3, in low cycle fatigue regime, LB-DED Inconel 718 specimens possessed better fatigue resistance at room temperature as compared to the isothermal high temperature fatigue tests conducted at 650°C. On the other hand, in high cycle fatigue regime, the effect of test temperature was observed to be minimum.

On the other hand, in high cycle fatigue regime, the effect of test temperature was observed to be minimum. Such variation in the fatigue behavior in different fatigue regimes of LB-DED Inconel 718 may be attributed to the change in microstructure resulting from the elevated test temperature.At high cycle fatigue regime, the majority of fatigue life is spent for the initiation of crack, which is not affected by the microstructure in presence of defects such as pores [11]. As a result, in the high cycle fatigue regime effect of test temperature on the fatigue life may have been minimal. Whereas, in the low cycle fatigue regime, majority of the fatigue life is spent in the crack growth state, which can be significantly affected by the microstructural features [11, 12]. Hence, significant effect of test temperature may have been evident on the fatigue life of LB-DED Inconel 718 at higher values of strain amplitudes (i.e. low cycle fatigue).

Hence, significant effect of test temperature may have been evident on the fatigue life of LB-DED Inconel 718 at higher values of strain amplitudes (i.e. low cycle fatigue). It has been reported that brittle precipitates will be formed at elevated temperatures on the grain boundaries, which exacerbate the fatigue resistance of material at elevated temperatures [13].5871.25LB-OED lnconel 718 (26° C)~ 0 -N --w <I ai "O ::, :t: a. E <( C: - ... CV u,10.750.50.25•AA... • i l • ....LB-OED lnconel 718 (650° C)A .,..' ' . A ~' '0 1.E+021.E+031.E+041.E+051.E+061.E+07Reversals to Failure, 2NtFigure 3: Comparison of fatigue life obtained for LB-DED Inconel 718 from the fatigue tests conducted at room temperature and at 650° CFigure 4 shows the fracture surface of the specimen subjected to 0.1 mm/mm strain amplitude at the 650° C.

To further understand the failure mechanism under isothermal high temperature fatigue loading, the fracture surface shown in Fig. 4 was compared with the one at obtained from room temperature test adopted from [7]. For the test conducted at the room temperature, the fracture surface was seen to be somewhat flat with the crack initiating from a metallic inclusion located at the surface of the specimen as shown in Fig. 4(a) and (b), which is presented in detail in [7]. On the other hand, for the test carried out at elevated temperature, the fracture surface was observed to be more irregular indicating a more torturous crack path, which may have resulted from the change in the microstructure during the test at elevated temperature. In addition, upon closer analysis, secondary cracks were also seen on the fracture surface of the specimen indicated by the red dashed arrow shown in the zoomed-in images in Fig.4.

In addition, upon closer analysis, secondary cracks were also seen on the fracture surface of the specimen indicated by the red dashed arrow shown in the zoomed-in images in Fig.4. Presence of such secondary cracks may be attributed to the formation of brittle precipitates at high temperature, which consequently affects the resulting isothermal high temperature fatigue lives of LB-DED Inconel 718.588Figure 4: Fracture surface of LB-DED Inconel 718 subjected to 0.1 mm/mm strain amplitude at 650ºC.ConclusionsIn this study, the fatigue behavior of Inconel 718 fabricated using a laser beam directed energy deposition (LB-DED) process was investigated under uniaxial strain-controlled fatigue tests at the elevated temperature of 650°C. Based on the results obtained from the experiments and fractography analysis, following conclusions are drawn:1. Testing temperature was seen to significantly affect the fatigue resistance of LB-DED Inconel 718 specimens.

Testing temperature was seen to significantly affect the fatigue resistance of LB-DED Inconel 718 specimens. Under identical value of strain amplitude, higher fatigue life was observed for the tests conducted at the room temperature.2. Larger variation in the fatigue life data was observed at low cycle fatigue regime, in which the majority of the fatigue life is governed by crack growth stage, when compared to the high cycle fatigue regime, in which the majority of the fatigue life is spent in crack initiation stage. Such a result can be attributed to the change in the microstructure resulting from the high test temperature, which primarily affects the crack growth behavior and not the crack initiation behavior.3.

Fractography analysis of specimen tested at high temperature revealed secondary cracks on the fracture surface, which may indicate the formation of brittle precipitates as a result of higher test temperature that can have a detrimental effect on the fatigue lives of LB- DED Inconel 718 specimens.589AcknowledgementsThe experiments were conducted when the authors were at Mississippi State University.References[1] F. Liu, X. Lin, C. Huang, M. Song, G. Yang, J. Chen, and W. Huang, The effect of laser scanning path on microstructures and mechanical properties of laser solid formed nickel- base superalloy Inconel 718, Journal of Alloys and Compounds, 509 (2011) 4505-4509.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFatigue Behavior of LB-PBF Ti-6Al-4V Parts Under Mean Stress and Variable Amplitude Loading ConditionsPatricio Carrion, Nima Shamsaei*Department of Mechanical Engineering, Auburn University, Auburn, AL 36849, USANational Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849, USACorresponding author: shamsaei@auburn.eduAbstractAdditively manufactured components are intended for use in load bearing applications, which are often accompanied by fluctuating external loadings. Therefore, understanding the fatigue behavior of AM materials under variable amplitude loadings is necessary for ensuring reliable in-service component performance.

Therefore, understanding the fatigue behavior of AM materials under variable amplitude loadings is necessary for ensuring reliable in-service component performance. This research focuses on the fatigue behavior of Ti- 6Al-4V, fabricated via laser beam powder bed fusion process, under mean stress and variable amplitude loadings. Mean stress effects are investigated under strain-controlled constant amplitude loading under fully-reversed, R(cid:304) = -1, and tension-release, R(cid:304) = 0, strain ratios. The generated data is used to compare two mean stress models including Morrow and Smith-Watson-Topper. Variable amplitude loading conditions include fully reversed high-low and low-high loading to investigate load sequence and history effects.

Variable amplitude loading conditions include fully reversed high-low and low-high loading to investigate load sequence and history effects. Finally, cumulative fatigue damage and life predictions are made using the linear damage accumulation model (i.e., Miner’s rule).Keywords: Additive manufacturing, Mean stress, Variable amplitude, Miner’s Rule, Titanium.IntroductionIn recent years, additive manufacturing (AM) technology has come to the spotlight as a promising alternative to conventional subtractive manufacturing methods. AM allows industries to fabricate far more complex parts directly from a CAD model by melting and joining the raw material layer-by-layer [1,2]. Laser beam powder bed fusion (LB-PBF) is among one of the most commonly used AM methods, which involves spreading a thin layer of metallic powder above the built platform and melting the powder to create a solid layer of the part [1,3,4].

Although innovative, parts fabricated via LB-PBF or other AM methods, are negatively affected by inherent internal and surface defects (e.g., porosity and as-built surface roughness). These types of defects have been identified as the most detrimental to fatigue performance and overall mechanical properties [2–13]. Post-processing methods, such as machining to remove the rough surfaces, and hot isostatic pressing (HIP) to close internal defects, have been proposed as feasible solutions to improve mechanical properties and performance of AM parts [2–5]. Unfortunately, post-646processing techniques are still not able to eliminate all of the defects and improve material properties to match those of the wrought counterparts.

Unfortunately, post-646processing techniques are still not able to eliminate all of the defects and improve material properties to match those of the wrought counterparts. Therefore, a current major concern is whether AM parts with minimum post-processing treatment (e.g., only stress relieving) can be integrated into load-bearing applications as a viable alternative to conventionally manufactured parts.Under realistic service conditions, AM parts will be subjected to random asymmetric fluctuating loads; therefore, the load history and its influence on the fatigue behavior with zero and non-zero mean stresses must be considered. In recent years, only a limited number of studies have attempted to characterize the fatigue behavior of AM parts under more realistic loadings [12,14–16]. The low and high cycle fatigue behavior of LB-PBF Ti-6Al-4V ELI (extra low interstitials) under different stress ratios (Rσ = -3, -1 and 0.1) was investigated in [14].

The low and high cycle fatigue behavior of LB-PBF Ti-6Al-4V ELI (extra low interstitials) under different stress ratios (Rσ = -3, -1 and 0.1) was investigated in [14]. The study concluded that the low cycle fatigue behavior is appropriately represented by the Coffin-Manson (C-M) equation, and that mean stress mainly affects fatigue behavior in the high cycle regime. Moreover, [15] focused on the variable amplitude fatigue behavior of rough as-built Ti-6Al-4V manufactured via LB-PBF and electron beam powder bed fusion (EB-PBF) process. In the study, it was demonstrated that conventional cumulative damage approaches could offer acceptable fatigue life predictions for parts subjected to fully random loading. To evaluate the load history and load sequence effects, the damage accumulated in a material under cyclic loading must be quantified.

To evaluate the load history and load sequence effects, the damage accumulated in a material under cyclic loading must be quantified. Therefore, an appropriate fatigue damage parameter (e.g., cycles to failure, crack size, stress response, etc.) must be utilized for fatigue life estimation using a suitable cumulative damage rule.The previously mentioned studies demonstrated the existence of a knowledge gap on the prospect of integrating AM parts directly from fabrication (i.e., minimum post-processing) to load- bearing applications. Therefore, as an effort to enhance the understanding of the fatigue behavior of AM parts, this research investigates as-built/annealed LB-PBF Ti-6Al-4V parts under mean stress and variable amplitude loadings. Mean stress effects are investigated using two strain ratios (R(cid:304) = -1, 0). Cumulative damage, load history, and load sequence effects were studied through fully reversed high-low and low-high loading.

Cumulative damage, load history, and load sequence effects were studied through fully reversed high-low and low-high loading. The generated constant amplitude data is used to evaluate the ability of Morrow, and Smith-Watson-Topper (SWT) mean stress correction models to predict fatigue life. Finally, the applicability of the linear damage rule (LDR) for fatigue life prediction of tests under high-low and low-high loading is evaluated.Experimental ProcedureSpecimen fabrication and material characterization Gas atomized Ti-6Al-4V Grade 23 metal powder with a mean particle size of ~30 μm manufactured by LPW Technology was selected for this study. The metal powder was used and was previously recycled. A brief explanation of the recycling/reconditioning procedure can be found in [11]. Round uniform gage section fatigue specimens were manufactured vertically using an EOS M290.

A brief explanation of the recycling/reconditioning procedure can be found in [11]. Round uniform gage section fatigue specimens were manufactured vertically using an EOS M290. The manufacturer recommended performance parameters used include 280 W laser power, 1200 mm/s scan speed, 100 mm stripe width, 140 μm hatch spacing, 67° hatch rotation, and a layer thickness of 30 μm. The uniform gage section had a 3.75 mm diameter and a 15 mm647length; these dimensions were based on ASTM E606 standard [17]. Post-manufacture, the specimens were stress relieved at 704 °C in an argon atmosphere for 1 hour, then furnace cooled to room temperature. Since only stress relief was performed, the specimens are considered to be in the “annealed as-built” condition.

Since only stress relief was performed, the specimens are considered to be in the “annealed as-built” condition. Microstructural characteristics of LB-PBF Ti-6Al-4V were investigated by dissecting the uniform gage section of a virgin specimen into two cross-sections: one cross-section perpendicular to the build direction, and one along the build direction. After dissection, the cross-sections were polished, etched using Kroll’s reagent, and a digital microscope was used to capture microstructure images.Test procedures Monotonic tensile and fatigue tests were performed using an MTS 810 servohydraulic system with a 100 kN load capacity. All tests were strain-controlled using an extensometer with a 10 mm gage length. Monotonic tensile tests (2 total) were first conducted at 0.001 s-1 strain rate until 0.045 mm/mm was reached, then switched to displacement control until specimen failure. The displacement rate used was 0.015 mm/s.

The displacement rate used was 0.015 mm/s. All fatigue tests were conducted following ASTM E606 standard [17] and using a tapered sinusoidal wave signal. For constant amplitude tests, fully reversed (R(cid:304) = -1) and pulsating (R(cid:304) = 0) loading was used. Variable amplitude tests include high- low and low-high two-level block loading. Strain amplitude, εa, values used in this study ranged between 0.001 - 0.006 mm/mm. To verify fatigue results, duplicate tests were conducted for each constant amplitude loading condition (e.g., R(cid:304) = -1 at different εa values), while for block loading (i.e., high-low and low-high at different εa values), three tests were conducted per condition. Experiments that reached over 4E6 reversals were determined to be a runout.

Experiments that reached over 4E6 reversals were determined to be a runout. Furthermore, strain rate effects on the fatigue behavior were attenuated by adjusting the test frequency in order to maintain a nearly constant average strain rate for all tests.Results and DiscussionMicrostructure Cross-sectional views of the microstructure of LB-PBF Ti-6Al-4V perpendicular to and along the build direction are presented in Fig. 1(a) and (b) respectively. It can be seen from the figures that the microstructure has fine acicular α(cid:397) grains surrounded by prior-β grain boundaries. From Fig. 1(a), it was also observed that the grain morphology is affected by the laser scan pattern and hatch distance (~140 μ m), thus resulting in equiaxed grains. However, epitaxial grain growth is observed in Fig. 1(b), which can be attributed to the high heat flow from the printed layer towards the relatively colder substrate, resulting in columnar grains oriented parallel to the build direction.

1(b), which can be attributed to the high heat flow from the printed layer towards the relatively colder substrate, resulting in columnar grains oriented parallel to the build direction. Similar results have been observed and extensively discussed in the literature for parts manufactured via LB-PBF process [5–9,11,10].648n o i t c e r i d d200 μμml iu B B200 μμm(a)(b)Fig. 1. Cross-sectional views of the microstructure of as-built LB-PBF Ti-6Al-4V (a) perpendicular to, and (b) along the build direction.Monotonic tensile behavior The monotonic stress-strain response of two LB-PBF Ti-6Al-4V specimens is presented in Fig. 2. As previously mentioned, the strain was recorded up to 0.0045 mm/mm, hence stress-strain curves end with arrows to indicate that the test continued in displacement control mode until specimen failure.

From the stress-strain curves, it can be seen that after yielding the material exhibits almost no work hardening; as a result, the yield stress value is relatively similar to the ultimate tensile stress. This behavior could be explained by the high yield strength of the martensitic α(cid:397) microstructure delaying the onset of permanent plastic deformation, but once plastic deformation occurs, the orientation of the retained prior-β grain boundaries facilitate plastic deformation as slip systems on preferred planes are activated (increasing ductility) without further increase in strength [6,18]. Moreover, due to the martensitic α(cid:397) microstructure found in LB-PBF Ti-6Al-4V, the yield and ultimate tensile strength are usually higher than wrought Ti-6Al-4V, which commonly has an equiaxed microstructure [4,6,10,12,19]. Monotonic tensile results include a yield (0.2 % offset) and ultimate tensile stress of 1040 and 1080 MPa, respectively, and 17%649elongation at failure.

The horizontal arrows indicate runout data, which are fatigue tests where no failure was observed after 4E6 reversals; the three runouts obtained are indicated by the “x3”. As seen in the figure, the strain-life plots present very low scatter within each loading condition (i.e., Rε = -1 and 0), with the exception of 0.001 mm/mm strain amplitude at Rε = -1. The as-built surface roughness condition (similar for all fabricated specimens with an average Ra = 16 μm) could be attributed to the low scatter in the low cycle fatigue regime as it allows failure to initiate from surface micro- notches rather than sub-surface defects, which are randomly located within the part [3,11]. In the high cycle regime, materials become more sensitivity to defects [3,20], hence, the wide scatter at 0.001 mm/mm strain amplitude may be attributed to the randomly located sub-surface defects also providing an opportunity for crack initiation and propagation.

Moreover, it can be seen from this figure that fatigue life (in both low and high cycle fatigue regime) is reduced for tests under Rε = 0 loading due to an increase in tensile mean stress and lack of stress relaxation, which can be attributed to the relatively low ductility of the material [19]. Therefore, contrary to the fatigue behavior of wrought Ti-6Al-4V which has been found not to be sensitive to mean stress under Rε = 0 loading in the low cycle fatigue regime [19], it can be concluded that as-built LB-PBF Ti-6Al- 4V is sensitive to mean stress.650In order to implement the C-M model for strain-life characterization, both R(cid:304) = -1 data sets were used. Since the elastic strain amplitude, (cid:304)a,e, was similar to the (cid:304)a, and negligible plastic strain amplitude, (cid:304)a,p, was observed, only the elastic part of the C-M model was needed and given by Eq. (1).

(1). The fatigue strength coefficient, (cid:305)(cid:10)f, and fatigue strength exponent, b, were obtained through linear regression of the (cid:304)a,e versus 2Nf.(cid:2013)(cid:3028) = (cid:2013)(cid:3028),(cid:3032) =(cid:4593) (cid:2026)(cid:3033) (cid:1831)(cid:3435)2(cid:1840)(cid:3033)(cid:3439)(cid:3029)Furthermore, as can be seen in Fig. 3, for the range of strain amplitudes used, the elastic portion of the C-M model accurately represents the fully reversed fatigue behavior of LB-PBF Ti-6Al-4V fabricated vertically and in the as-built/annealed condition. This is expected as LB-PBF Ti-6Al- 4V is considered a relatively brittle material thus the presence of plastic deformation is minimal.)m m m m/0.0070.0060.005-••R = -1, PS R = -1 ε R = 0 ε C-Mε(a (cid:304), e d u t i l p m A n i a r t S0.0040.0030.0020.001••◊ -x30.0001.E+31.E+41.E+51.E+61.E+7Reversals to Failure, 2NfFig. 3.

3. Strain-life plot including fully reversed data from powder recycling study (R(cid:304) = -1, PS) [11], fully reversed (R(cid:304) = -1) and pulsating (R(cid:304) = 0) data from this study, and the Coffin-Manson (C-M) fit.Since the stress-strain response of LB-PBF Ti-6Al-4V is fully elastic, the analysis of the fatigue behavior under mean stress was conducted using stress-based correction models, these include Morrow, and Smith-Watson-Topper (SWT) [20–22]. Both models are presented in Eq. (2), and (3), respectively, where σa is the stress amplitude, σmax is the maximum stress, and σm is the mean stress. The purpose of the models is to provide an equivalent fully reversed stress amplitude, σeq, that yields a similar result in fatigue life as those with non-zero mean stress; then a modified version of Eq. (1) (to stress amplitude instead of strain amplitude) is used to predict fatigue life and presented in Eq. (4) [19–21,23,24].

4(a) and (b), respectively. In the figures, the solid line represents a 1:1 correlation between predicted and experimental fatigue lives and the dashed lines represent scatter bands with a factor of two. Horizontal arrows indicate data points for which the predicted fatigue life was finite while the experimental life was a runout. From Fig. 4(a), it can be seen that Morrow significantly underestimates the effects of mean stress, resulting in non-conservative fatigue life predictions and poor correlation (R2 = 0.86) between experimental and predicted fatigue lives. On the contrary, it has been shown that Morrow is able to offer better fatigue life predictions for wrought Ti-6Al-4V [19]. The SWT model, shown in Fig. 4(b), offers fatigue life predictions well within the scatter bands of two, effectively reducing scatter and providing a better correlation (R2 = 0.99).

The SWT model, shown in Fig. 4(b), offers fatigue life predictions well within the scatter bands of two, effectively reducing scatter and providing a better correlation (R2 = 0.99). Acceptable correlations obtained with the SWT approach may be explained by the ability of the model to consider that for high strength materials, with low plasticity, failure is often driven by the maximum tensile stress [12,14,20].1E+71E+7R2 = 0.86R2 = 0.99w o r r o M N 2, e f i Ld e t c i d e r P1E+61E+51E+4R = -1, PS R = -1 ε R = 0 εεT W S N 2, e f i Ld e t c i d e r P1E+61E+51E+4R = -1, PS R = -1 ε R = 0 εε1E+31E+31E+31E+41E+5 Experimental Life, 2Nf1E+61E+71E+31E+41E+5 Experimental Life, 2Nf1E+61E+7(a)(b)Fig. 4. Comparison of fatigue life predictions using (a) Morrow, and (b) Smith-Watson-Topper (SWT) mean stress models with experimental results.Cumulative damage Presented in Fig.

4. Comparison of fatigue life predictions using (a) Morrow, and (b) Smith-Watson-Topper (SWT) mean stress models with experimental results.Cumulative damage Presented in Fig. 5 is a semi-log plot of the strain amplitude, (cid:304)a, versus cycles to failure, Nf, for R(cid:304) = -1 and (R(cid:304) = -1, PS) [11] constant amplitude tests, as well as the second loading block652from high-low and low-high tests. The horizontal arrows indicate data runouts, the “x3” indicates runouts for R(cid:304) = -1, and the “x2” indicates the runouts from high-low loading. From the figure, it can be observed that the load sequence does affect the fatigue behavior of LB-PBF Ti-6Al-4V. Low-high tests that were preloaded with 0.003 mm/mm strain amplitude result in shorter fatigue lives when compared to the constant amplitude equivalents (0.006 and 0.005 mm/mm strain amplitude).

Low-high tests that were preloaded with 0.003 mm/mm strain amplitude result in shorter fatigue lives when compared to the constant amplitude equivalents (0.006 and 0.005 mm/mm strain amplitude). The decreased fatigue life for the first load sequence may be explained by surface defects, because of the as-built surface condition, expediting crack initiation [3,20,25]. As a result, only crack propagation takes place during the second load sequence. It is important to mention that the observed behavior is not typical to conventional engineering materials [25,26]. For example, wrought Ti-6AL-4V subjected to low-high loading results in longer fatigue life for the second loading block due to material exercising during the first block [27,28]. Tests preloaded with 0.001 mm/mm strain amplitude, followed by 0.003 mm/mm strain amplitude loading (showed by the enhanced view in Fig.

Tests preloaded with 0.001 mm/mm strain amplitude, followed by 0.003 mm/mm strain amplitude loading (showed by the enhanced view in Fig. 5), resulted in equal or longer fatigue lives; which may be justified by the first sequence being small enough that crack initiation is not affected [20,25]. The trends observed during high-low loading may also be justified by the same surface roughness phenomena. During the high load, crack initiation and propagation are expedited due to surface defects, and during the low load, the cracks only have to propagate until failure is reached, resulting in fatigue life reduction.

Finally, it was also observed that if the second sequence of high-low loading is at 0.001 mm/mm strain amplitude, the resultant stresses are not enough for crack propagation.0.007)m m m m/(0.0060.005...◊◊+ x2R = -1, PS R = -1 Preloaded with High Preloaded with Lowε◊ <> Aεa (cid:304), e d u t i l p m A n i a r t S0.0040.0030.002~0.001◊O & <>-+ IJ.. + + x3 x20.0001.E+21.E+31.E+41.E+51.E+61.E+7Cycles to Failure, NfFig. 5. Strain-life comparison of fatigue lives obtained from the second load block (i.e., low for high-low, and high for low-high) with fully reversed data generated in this study (R(cid:304) = -1) and a powder recycling study (R(cid:304) = - 1, PS) [11].The induced damage from two-level block loading, i.e., high-low and low-high, on LB- PBF Ti-6Al-4V was calculated using Miner’s rule [20,25,29] and the strain-life C-M model (see Fig. 3).

3). Fatigue life predictions for the second loading block (i.e., remaining fatigue life) using LDR are presented versus the experimentally obtained fatigue lives in Fig. 6. In the figure, the solid line represents a 1:1 correlation between predicted and experimental fatigue lives, the dashed lines are scatter bands with a factor of two, and “x2” indicates that there are two data points. As can be seen in Fig. 6, the LDR offers rather conservative fatigue life predictions since all data653points fall below the 1:1 line. Furthermore, an R2 = 0.96 indicates that the LDR is able to appropriately correlate the cumulative damage behavior of LB-PBF Ti-6Al-4V.

Furthermore, an R2 = 0.96 indicates that the LDR is able to appropriately correlate the cumulative damage behavior of LB-PBF Ti-6Al-4V. The successful correlation of the aforementioned damage behavior using the LDR may be elucidated by the pre- existing surface defects which, regardless of loading (H-L and L-H), are able to propagate during both load sequences.1E+5R2 = 0.96R D L, 2N, e f i L g n i n i a m e R d e t c i d e r P1E+41E+3x2•H-L L-H1E+21E+21E+31E+41E+5Experimental Remaining Life, N2Fig. 6. Miner’s rule fatigue life predictions for the second loading block (i.e., remaining life), from high-low (H- L) and low-high (L-H) compared with experimental fatigue lives.ConclusionsThe microstructural characteristics, monotonic tensile deformation behavior, and fatigue behavior of as-built/annealed LB-PBF Ti-6Al-4V were investigated.

The Morrow and Smith- Watson-Topper (SWT) mean stress correction models were employed to correlate R(cid:304) = 0 fatigue data. The linear damage rule (LDR) in conjunction with the Coffin-Manson (C-M) fit was used to capture the cumulative damage for high-low and low-high tests. Based on the experimental results and discussions, the following conclusions are drawn:1. The microstructure characteristics and mechanical properties of LB-PBF Ti-6Al-4V did not present any significant differences from results found in the literature for vertically fabricated LB-PBF parts. Moreover, the monotonic stress-strain response exhibited similar values for yield and ultimate tensile stress due to negligible work hardening, which was attributed to the martensitic α(cid:397) microstructure combined with the prior-β grain boundaries.2. LB-PBF Ti-6Al-4V is sensitive to mean stress across both low and high cycle fatigue regimes.3.

LB-PBF Ti-6Al-4V is sensitive to mean stress across both low and high cycle fatigue regimes.3. The Smith-Watson-Topper mean stress correction model was able to provide appropriate fatigue life predictions while Morrow offered non-conservative predictions due to failure being driven mainly by the maximum tensile stress.4. Relatively small load sequence effects were observed for LB-PBF Ti-6Al-4V under high-low and low-high loading. These results were justified by the as-built surface654condition enabling cracks to initiate/propagate during both blocks of the load sequences.5. The LDR successfully correlated the cumulative damage behavior and accounted for load sequence effects for LB-PBF Ti-6Al-4V and provided conservative fatigue life predictions.AcknowledgmentsThis research was partially funded by the National Science Foundation (NSF) under grantNo.1657195References[1] Frazier WE. Metal Additive Manufacturing: A Review. J Mater Eng Perform 2014;23:1917–28.

TIGUE BEHAVIOR OF SELECTIVE LASER MELTED 17-4 PH STAINLESS STEELAref Yadollahi1, Nima Shamsaei1,*, Scott M. Thompson1, Alaa Elwany2, Linkan Bian3 & Mohamad Mahmoudi21Department of Mechanical Engineering, Mississippi State University, Mississippi State, MS 39762 1Center for Advanced Vehicular Systems (CAVS), Mississippi State University, Mississippi State, MS 39762 2Department of Industrial and Systems Engineering, Texas A & M University, College Station, TX 3Department of Industrial and Systems Engineering, Mississippi State University, Mississippi State, MS 39762Corresponding author: Email: shamsaei@me.msstate.edu Phone: (662) 325 2364AbstractIn this investigation, fully-reversed strain-controlled fatigue tests were conducted on Selective Laser Melted (SLM) 17-4 PH stainless steel (SS). Cylindrical 17-4 PH rods were fabricated vertically-upward using optimized process parameters to ensure a dense product.

Cylindrical 17-4 PH rods were fabricated vertically-upward using optimized process parameters to ensure a dense product. Post- fabrication heat treatments (solution annealing and aging) were applied on half of the as-built samples. Fatigue behavior and tensile properties of the as-built and heat treated samples were investigated and compared with available data from the literature. The microstructure analysis and fractography were performed to discern the failure initiation sites, crack propagation path, and fracture surface morphology. Fatigue lives of SLM 17-4 PH SS specimens were found to be significantly shorter than their wrought counterparts. It was also found that heat treatment hardens the SLM 17-4 PH SS specimens while also shortens their fatigue life in the high cycle regime.

It was also found that heat treatment hardens the SLM 17-4 PH SS specimens while also shortens their fatigue life in the high cycle regime. The presence of defects, which serve as crack initiation sites, and more sensitivity of heat treated specimens to impurities, due to higher hardness, were the main reasons for these observations.IntroductionDuring the past decade, increasing attention has been drawn to additive manufacturing (AM), a unique manufacturing process in which a part is built up layer-by-layer in contrast to traditional, ‘subtractive’ metalworking processes. Additive manufacturing is currently demonstrating its proficiency in fabricating components with sophisticated geometries targeted towards biomedical, aerospace, and defense industries [1]. Of the AM techniques available, Selective Laser Melting (SLM), which is a laser based powder bed system, provides a means to obtain relatively fine surface quality, high precision, powder efficiency, etc.

For a typical SLM process, a thin layer of metallic powder is placed on the substrate and a focused laser follows a pattern over the surface, melting powder particles to form a part layer by layer. The build process is conducted in an inert atmosphere (e.g. argon).Fabricating AM components with microstructural and mechanical properties comparable to those produced by conventional manufacturing techniques is still a major challenge. This is due to the many involved process parameters (such as: laser power, beam travel speed, layer thickness, scanning strategy, part building orientation) providing for complex thermal histories that then lead to the formation of a non-homogeneous, anisotropic microstructure with inevitable porosity [2].

In general, static mechanical properties (such as tensile, compression, hardness, etc.) of AM parts are comparable or even higher than those of wrought and cast materials due to relatively high cooling rates experienced during fabrication which favors the production of finer microstructures. However, lower fatigue strength has been reported for AM parts due to detrimental effects caused by defects such as pores and inclusions - inherent consequences of many powder-based AM techniques. Thus, finding effective optimization and control mechanisms to fabricate components via AM with superior mechanical properties is of interest [3].Among the many steels that have been processed with SLM, precipitation hardened (PH) steels are more attractive due to their weldability and austenitic/martensitic microstructure [4]. These alloys have been used for structures in nuclear, aerospace, marine, naval, and chemical industries [5,6].

These alloys have been used for structures in nuclear, aerospace, marine, naval, and chemical industries [5,6]. Stainless steel 17-4 PH is one of the most popular alloys in this category due to its relatively high tensile/impact strength, fracture toughness and corrosion resistance at typical service temperatures below 300 ºC [7].Microstructural and mechanical properties of SLM 17-4 PH SS have been reported and discussed in the literature [8–11], and the high cycle fatigue behavior of both additively- and conventionally-manufactured 17-4 PH SS under stress-control tests has been investigated extensively [5,10,12]. However, to the best of authors’ knowledge, there is no available report on the strain-controlled low cycle fatigue behavior of this material. Therefore, the aim of this study is to investigate strain-controlled fatigue behavior of SLM 17-4 PH SS.

Therefore, the aim of this study is to investigate strain-controlled fatigue behavior of SLM 17-4 PH SS. The effect of an appropriate heat treatment (solution annealing and aging) on fully-reversed fatigue strength of SLM 17-4 PH SS is also studied. Fatigue fracture surfaces are also examined using Scanning Electron Microscopy (SEM) to capture the crack initiation site, crack propagation and failure mechanisms involved.Experimental ProceduresCommercially-procured, gas-atomized 17-4 PH SS powder (Phenix Systems), with 80 percent of its size distribution falling below 22 µm, was utilized for fabricating rods via a SLM machine (ProX™ 100). Process parameters (i.e. laser power, scanning speed, layer thickness, and hatching pitch) were selected to obtain an acceptable level of final part density using a sequential experimental methodology [13]. Utilized process parameters (i.e. laser power, scanning speed, layer thickness, and hatching pitch) are summarized in Table 1.Table 1.

Utilized process parameters (i.e. laser power, scanning speed, layer thickness, and hatching pitch) are summarized in Table 1.Table 1. Utilized SLM process parameters for fabricating 17-4 PH SS rods.Laser Power (W) Scanning Speed (mm/s) Hatching Pitch (μm) Layer Thickness (μm)483005030Cylindrical rods, with 8 mm diameter and 75 mm height, were fabricated upward in a vertical orientation on a base plate. After fabrication, the specimens were removed from the base plate via Electrical Discharge Machining (EDM). In order to investigate the effect of heat treatment, half of the as-built samples were subjected to solution annealing for 30 min at approximately 1040 °C followed by air cooling (AC) to room temperature (Condition A) and then, these same samples underwent precipitation hardening for 1 hour at 482°C followed by AC to room temperature (Condition H900).

It has been reported that when subjecting as-built SLM 17-4 PH SS samples to direct aging, age hardening will not occur due to the presence of dual phase microstructure (i.e. martensitic and austenitic phases) [8]. Therefore, solution annealing was performed on the as-built samples before peak-aging. Heat treated and non-heat-treated tensile and fatigue tests specimens were machined from the SLM samples to the configuration and dimensions in conformance to the ASTM E466 standard [14], as shown in Fig. 1. After machining, the specimens were polished to a roughness of Ra < 0.7 µm in order to minimize surface defects. The polished specimen roughness was determined using a non-contact (i.e. laser reflectance) measurement system (Talysurf).Figure 1.

The polished specimen roughness was determined using a non-contact (i.e. laser reflectance) measurement system (Talysurf).Figure 1. Dimensions of fatigue and tensile specimens according to ASTM E 606 [14].Tension experiments were conducted at room temperature using an Instron 5882 servo- hydraulic testing machine with a ±100 kN maximum loading capacity at a 0.001/s nominal strain rate. Fully-reversed strain-controlled fatigue tests, with a strain ratio of R = -1, were performed with an MTS Tabletop 858 machine under various strain amplitudes. Fatigue experiments were performed in accordance with ASTM E606 under a sinusoidal loading waveform until failure occurred or 106 cycles was reached, in which the test was considered to be a run-out. Fatigue test frequencies were adjusted according to strain amplitude. One specimen from each set was mounted and etched for microstructural characterization using optical microscopy (OM) (ZEISS Axiovert 200).

One specimen from each set was mounted and etched for microstructural characterization using optical microscopy (OM) (ZEISS Axiovert 200). Fracture surfaces of fatigued specimens were Au/Pd sputter-coated and examined using scanning electron microscopy (SEM) (FEG SEM Zeiss SUPRA™ 40) to determine crack initiation site and crack propagation characteristics.Results and DiscussionsMicrostructureThe microstructure of SLM parts depends primarily on the temperature gradients and cooling rates (i.e. thermal history) experienced during the manufacturing process. The utilized process parameters will directly impact the heat transfer into the powder bed, and this will then affect the generated melt pool size, heat affected zone (HAZ), and cooling rate [11].

The utilized process parameters will directly impact the heat transfer into the powder bed, and this will then affect the generated melt pool size, heat affected zone (HAZ), and cooling rate [11]. The heat-dependent, encumbered microstructure of SLM parts is of importance as it will impact their mechanical properties.Optical micrographs of the longitudinal cross-section of SLM 17-4 PH SS samples, in the as-built and heat treated conditions, are shown in Figs. 2(a) and (b), respectively. The microstructural examination revealed the presence of voids; including pores due to entrapped gas and un-melted regions between layers due to insufficient fusion combined with low laser penetration depth. Such defects can impact the mechanical properties of the material, especially elongation to failure and fatigue behavior. It was also found that post-manufacturing heat treatment has a substantial impact on the microstructure.

It was also found that post-manufacturing heat treatment has a substantial impact on the microstructure. After solution annealing and peak- aging, the microstructure transformed to a fully recrystallized structure, as presented in Fig. 2(b). Precipitation at grain boundaries during the peak-aging treatment can be clearly seen in Fig. 2(b), since these regions were easier to be etched. Heat treatment, as expected, homogenized the microstructure as compared to the non-heat-treated samples, as shown in Fig. 2(a). In addition, it can be observed that heat treatment eradicates interfacial regions between deposited layers.(a)(b)Figure 2. Longitudinal cross-section images of SLM 17-4 PH SS in (a) as-built and (b) heat treated conditions.Tensile PropertiesTensile properties of the heat treated and non-heat-treated SLM 17-4 PH SS are presented in Fig. 3.

3. As shown, the yield and ultimate tensile strengths are significantly increased after heat treatment (solution annealing and aging), but elongation to failure is reduced. This behavior is due to precipitation hardening of the chromium-nickel-copper constituents which prevents dislocation movement causing metal to resist deformation and become harder [15]. The monotonic tension results reveal that the yield and ultimate tensile strengths of SLM 17-4 PH SS in as-built condition are lower than those of wrought material in H900 condition (yield stress: 1,170 MPa, ultimate tensile stress: 1,310 MPa) [16]. This is due to the fact that the SLM 17-4 PH SS samples are not completely martensitic, but contain retained austenite due to high solidification speeds during fabrication, which prevent the formation of the martensite phase in the as-built material, leading to a metastable austenitic microstructure.

The presence and amount of retained austenite may have a significant influence on the material’s strength, strain hardening and elongation to failure [6,8]. Therefore, a post-manufacturing heat treatment is necessary toenhance the mechanical properties of SLM 17-4 PH SS. The yield and ultimate tensile strengths of heat treated SLM 17-4 PH SS is comparable to those of conventionally-manufactured parts in the H900 condition. However, elongation to failure of heat treated specimens is noticeably lower than wrought materials in the H900 condition (10% elongation to failure). Lower ductility of SLM 17-4 PH SS is attributed to the presence of defects, such as pores and un-melted powder, resulting from entrapped gas and/or lack of fusion during fabrication.Figure 3. Engineering stress-strain curves of SLM 17-4 PH SS in as-built and heat treated conditions.Fatigue BehaviorFatigue is the main mechanical failure mode for most components and structures.

Engineering stress-strain curves of SLM 17-4 PH SS in as-built and heat treated conditions.Fatigue BehaviorFatigue is the main mechanical failure mode for most components and structures. Unlike failure due to sudden or monotonic loads, fatigue is a localized and microstructural-dependent failure mode resulting from cyclic loading over longer time scales. In fact, fatigue failure, as a result of cyclic loading, can occur due to a stress that is much lower than failure-causing monotonic stress. The three characterized stages of fatigue failure are: crack initiation (nucleation), crack propagation and final fracture – all corresponding to categorized regions of fatigue fracture surfaces. Generally, cracks tend to initiate from microstructural defects (i.e. pores and particles) and grain boundaries. However, in the case of AM/SLM parts, defects can be the main source of crack initiation [3].

pores and particles) and grain boundaries. However, in the case of AM/SLM parts, defects can be the main source of crack initiation [3]. These features may create local microscopic stresses larger than the yield strength, causing local plastic deformation and leading to fatigue crack initiation under cyclic loading.Fatigue life is typically quantified by the number of load cycles/fluctuations until failure, 𝑁f. Generally speaking, fatigue life that is less than 103 cycles is considered as low cycle fatigue (or short-life regime), fatigue life in the range of 103 < 𝑁f < 105 is classified as mid-cycle fatigue (or mid-life regime), and 𝑁f > 105 is categorized as high-cycle fatigue (or long-life regime). The most commonly used fatigue life models (to predict 𝑁f) are the nominal stress-life (𝑆 − 𝑁) model and the local strain-life (𝜀 − 𝑁) model. The stress-life method is traditionally used for the fatigue life analysis at the high cycle fatigue (HCF) regime.

The stress-life method is traditionally used for the fatigue life analysis at the high cycle fatigue (HCF) regime. This method relates the nominal stresses to local fatigue strengths. The strain-life method involves a more thoroughanalysis of the plastic deformation at localized regions, and consequently, can better characterize the fatigue behavior of a material, specifically in the presence of plastic deformations - making it the most appropriate for low cycle fatigue (LCF) applications.The results of uniaxial fully-reversed (R = -1) fatigue tests of SLM 17-4 PH SS, performed at various strain amplitudes, are illustrated as (𝜀 − 𝑁) curves in Fig. 4. For a more conservative estimation, the run-out specimens were not considered for curve fitting. The scatter in results was greatest for tests conducted in longer life regimes and least for tests conducted at shorter life regimes. This is due to the fact that defect size and location (i.e.

The scatter in results was greatest for tests conducted in longer life regimes and least for tests conducted at shorter life regimes. This is due to the fact that defect size and location (i.e. distance from surface) significantly affects the fatigue life in the HCF regime, where crack initiation constitutes a dominant part of total life. As shown in Fig. 4, the heat treated SLM specimens demonstrate higher fatigue strengths relative to their as-built counterparts in LCF, but the opposite trend occurs for HCF. The behavior of heat treated SLM specimens in HCF contrasts the expectation that the high cycle fatigue strength of steel typically increases with an increase in tensile strength [17].It is well known that crack initiation dominates total fatigue lifetime in HCF [18].

Therefore, it is expected that the heat treated SLM specimens exhibit more HCF life as compared to their as-built counterparts - since they possess higher strength and more resistance to dislocation movement within the matrix, both resulting in enhanced resistance to crack initiation. Moreover, coherent copper-rich precipitates in the matrix of the peak-aged (H900 condition) specimens are expected to provide more resistance to crack initiation. However, this is not the case for AM specimens that already contain defects as large as 100 m in some cases. Therefore, defect size, shape, and location may play a more dominant role in HCF life of AM/SLM parts. In addition, heat treated specimens are harder, and therefore, more sensitive to impurities, specifically in HCF regime.Figure 4.

In addition, heat treated specimens are harder, and therefore, more sensitive to impurities, specifically in HCF regime.Figure 4. Strain-life fatigue experimental data and curves for SLM 17-4 PH SS in both as-built and heat treated conditions.Maximum stress-life fatigue data of heat treated SLM 17-4 PH SS and their comparison with wrought material in the H1050 condition [19] is presented in Fig. 5(a). The fatigue stress amplitude versus life of SLM 17-4 PH SS (in the as-built condition) is also compared with reported performance data for as-built Direct Metal Laser Sintered (DMLS) 17-4 PH SS [10] in Figure 5(b). The fatigue data for wrought material (H1050 condition) reported in Fig. 5(a) belong to the flat specimens (KT = 1) tested under completely reversed axial loading (R = -1) [19]. The fatigue data reported in Figure 5(b) belong to DMLS horizontally-built specimens tested using a rotating bending testing machine (R = -1) [10].

The fatigue data reported in Figure 5(b) belong to DMLS horizontally-built specimens tested using a rotating bending testing machine (R = -1) [10]. The surface of each specimen was polished in both studies. Although general observations are warranted, there is inadvertent variance in fatigue data due to different testing modes, specimen types, load setups, fabrication processes, etc.; hence, it is difficult to make direct comparisons between results from this study and others in literature.(a)(b)Figure 5. Comparison of fatigue stress life data (a) for heat treated SLM 17-4 PH SS (this study) to wrought 17-4 PH SS in H1050 condition from [19] and (b) for as-built SLM 17-4 PH SS (this study) to DMLS 17-4 PH SS in as-built condition from [10].As shown in Figure 5(a), SLM 17-4 PH SS demonstrates significantly shorter fatigue lives as compared to the wrought material in the H1050 condition.

This is attributed to the presence of a large number of defects including pores, un-melted regions, and un-melted powder particles which serve as crack initiation sites. Moreover, the ultimate tensile property reported by Leybold [19] (1,414 MPa) is significantly higher than those measured for the heat treated SLM 17-4PH SS in this study. From Figure 5(b), it can be seen that the mid and high cycle fatigue strength of the as-built SLM 17-4 PH SS is also lower than the DMLS data reported by Sehrt and Witt [10]. This can be due to differences in (1) building orientation (i.e. vertical in this study versus horizontal in Sehrt and Witt [10]) and (2) utilized process parameters which affect the thermal histories during fabrication, and consequently, microstructure and phase proportion of the material.

Furthermore, a rotating bending test may be less severe than a uniaxial tensile fatigue tests (employed in this study), since the whole gage section is subjected to the maximum stress for the latter, while only portions of the gage surface experiences the maximum stress under rotating bending. Statistically, more impurities can serve as the crack initiation site in closed loop uniaxial fatigue machine as compared to the rotating bending test where less volume and impurities are exposed to maximum stress.FractographyThe analysis of the fatigue fracture surfaces of the heat treated and non-heat-treated SLM 17-4 PH SS specimens reveals that crack initiations occur at defects adjacent to specimens’ surface, including pores due to entrapped gas, and un-melted regions (or weak metallurgical bounding between layers) resulting from lack of fusion (or low penetration depth of laser).

This indicates that although machining and polishing reduces the surface roughness (due to partially- melted powder particles and the inconsistency of melt pool) and associated notch effects, remnant interior pores and sub-surface defects that are brought to the surface still play a dominant role in crack initiation. Scatter in fatigue life data have a direct relationship with size and location (i.e. distance from surface) of pores and un-melted regions. Larger un-melted regions and/or ones closer to the surface have more detrimental effects on fatigue strength, especially during HCF. For specimens that failed relatively early (in HCF), larger un-melted regions have been detected near the specimens surface.Figure 6(a) demonstrates a typical fatigue fracture surface of an as-built SLM specimen that failed in the long life regime (i.e. low strain amplitude). As shown in Fig.

low strain amplitude). As shown in Fig. 6(b), an un-melted region close to the specimen surface acts as a stress raiser and serves as a site for crack initiation. However, multiple crack initiation sites were found on the fracture surface of specimens tested in the short life regime, as depicted in Fig. 7(a). As can be seen in Figs. 7(b) and (c), for a specimen tested in LCF, cracks appear to have initiated from surface porosity and un-melted regions beneath the specimen surface. This is due to the fact that crack progression takes more time in LCF, which provides an opportunity for other cracks to initiate. In addition, the defects that act as crack initiation sites in LCF appear to have smaller sizes as compared to those found on the fracture surface of specimens tested in HCF, as can be seen in Figs. 7(b) and (c). This may indicate that the influence of defect size is less pronounced than its distance from the surface in short life regimes.

7(b) and (c). This may indicate that the influence of defect size is less pronounced than its distance from the surface in short life regimes. In other words, the fatigue specimens (both conditions) during the long life regime demonstrated internal (i.e. subsurface) crack initiation while the ones in short life regime mostly demonstrated surface crack initiation.Generally, three main regions that correspond to fatigue crack initiation, fatigue crack growth, and final fracture can be observed on all fatigue fracture surfaces. The fraction of areas of crack propagation and final fracture depend on the applied strain level. Lower applied strains (i.e. long life regimes) result in a large stable crack propagation area, as shown in Fig. 6(a). On the other hand, for higher strain levels (i.e. short life regimes), the crack will not be able to grow very long before final fracture, resulting in a smaller crack growth area, as seen in Fig. 7(a).

short life regimes), the crack will not be able to grow very long before final fracture, resulting in a smaller crack growth area, as seen in Fig. 7(a). Similar crack initiation site and morphology of fracture surface were also observed for heat treated fatigue specimens in both LCF and HCF. Voids and un-melted powder (attributed to entrapped gas in melt pool and lack of melting during fabrication, respectively) can be seen along the final fracture surface of specimens in both as-built and heat treated conditions.(a)(b)Figure 6. (a) An overall view of fracture surfaces for as-built SLM 17-4 PH SS at low strain 0.15% (Nf = 236,820) and (b) a magnified view of the crack initiation site indicating that fatigue cracks initiated from an un-melted region close to the surface.(b)(a)(c)Figure 7.

(a) An overall view of fracture surfaces for as-built SLM 17-4 PH SS at strain amplitude of 0.3% (Nf = 6,422), and a magnified view of the crack initiation sites including (b) pores beneath the specimen surface and (c) an un-melted region close to the surface.ConclusionsIn this study, the fatigue behavior of 17-4 PH stainless steel fabricated via SLM in both the as-built and heat treated (solution annealing and aging) conditions was investigated. A strain- controlled uniaxial test setup with a strain-ratio of R = -1 (fully-reversed) was employed. The following conclusions can be drawn based on the experimental results:1. Defects inherent to powder-based AM has a substantial effect on fatigue life, especially in long life regimes. Unexpectedly, HCF lives of heat treated SLM specimens were lower than their as-built counterparts. This may be explained by heat treatment hardening 17-4 PH SS, making the material more sensitive to impurities.2.

This may be explained by heat treatment hardening 17-4 PH SS, making the material more sensitive to impurities.2. Fatigue strength of heat treated SLM 17-4 PH SS was found to be significantly lower than wrought material in the H1050 condition. Aside from differences in including pores and un-melted microstructure, regions/powder, were the main contributors for the shorter fatigue life of SLM 17-4 PH SS. the presence of defects,3. The fatigue strengths of the investigated SLM 17-4 PH SS specimens were found to be lower than those measured from 17-4 PH SS made using a similar AM processing technique. This may be explained by differences in the employed fatigue test setup (i.e. rotating bending versus closed loop uniaxial) and/or the design (i.e. building orientation) and process parameters utilized.4.

rotating bending versus closed loop uniaxial) and/or the design (i.e. building orientation) and process parameters utilized.4. Fracture surfaces of fatigue specimens revealed that cracks were mostly initiated from un-melted regions that formed due to insufficient fusion, and pores, which may have formed via entrapped gas. Cracks initiated from large un-melted regions adjacent to the surface of specimens failed under high strain amplitude (i.e. long life regime). Multiple crack initiation sites on the surface were observed for specimens tested in shorter life regime. References[1]S.M. Thompson, L. Bian, N. Shamsaei, A. Yadollahi, An Overview of Direct Laser Deposition for Additive Manufacturing; Part I: Transport Phenomena, Modeling, and Diagnostics, Submitt. to Addit. Manuf.[2] A. Yadollahi, D. Seely, N. Shamsaei, S.M.

TIGUE CHARACTERIZATION OF 3D PRINTED ELASTOMER MATERIALJacob P. Moore and Christopher B. Williams Design, Research and Education for Additive Manufacturing Systems Laboratory Department of Mechanical Engineering Virginia Polytechnic Institute and State UniversityAbstractThe Objet PolyJet 3D Printing process provides the ability to print graded materials featuring both stiff and elastomeric polymers. This capability allows for a variety of new design possibilities for additive manufacturing such as living hinges, shock absorbing casings, and integrated gaskets. Such design features typically rely upon the ability of traditional elastomers to experience large and repeated strains without permanent deformation or damage. However, voids and other flaws inherent to many Additive Manufacturing (AM) processes can have a significant negative impact on the fatigue life of elastomeric AM materials.

However, voids and other flaws inherent to many Additive Manufacturing (AM) processes can have a significant negative impact on the fatigue life of elastomeric AM materials. In this paper, the authors seek to fill a gap in the literature by characterizing the fatigue life of a direct 3D printed elastomer, and the multi-material interface. Based on the results, the authors offer advice for improving fatigue life of printed elastomeric components.Keywords: Direct 3D Printing, Objet PolyJet, Elastomer, Fatigue1. Background and MotivationThe Objet PolyJet process is a direct 3D printing process (direct 3DP) in which layers of photopolymer are printed onto the build platform via inkjet print heads. During printing, UV lamps (located on either side of the print head) cure the recently deposited photopolymer.The direct 3DP process differs from the more common indirect 3DP processes in that the build material is being jetted, rather than jetting a binder into a bed of build material.

Because there is no powder bed in direct 3DP processes, multiple materials can be printed in the same build – similar to the manner in which multiple colors can be printed with traditional inkjet printers.Figure 1: An illustration of the Objet direct 3DP process [1]Objet has developed a number of different photopolymers with differing degrees of stiffness, strength, and transparency TangoBlackPlus a “rubber-like” photopolymer has the lowest Young’s Modulus (E) of any PolyJet material [1]. The ability to print elastomers such as TangoBlackPlus using the Objet PolyJet process, particularly the ability to print elastomers integrated with other stiffer materials, creates an opportunity for a variety of new design possibilities for additive manufacturing.

Potential applications of 3D printed elastomers include641integrated shock absorption components, integrated grips, living hinges, gaskets, and other flexible components that traditionally utilize the unique properties of elastomers. An example of a PolyJet part with integrated elastomer living hinges (the black material) is shown in Figure 2.One of the key characteristics of traditional elastomers is their ability to experience large repeated strains without permanent deformation or failure. With Objet labeling these new materials as “rubber-like,” many designers may expect the material to exhibit the same abilities of traditional elastomers. However, the flaws and voids caused by the layer-by-layer fabrication process in many AM processes [2] may significantly reduce the fatigue life of printed elastomers, since fatigue failure in elastomers is the result of the propagation of small tears present in the specimen due to manufacturing [3, 4].

As there is no data available on the fatigue characteristics of the TangoBlackPlus material, designers have no way to predict the lifetime of the elastomeric components of their product. This paper seeks to fill this gap in the literature by:Figure 2: A 3DP part with elastomer living hinges (shown in black) [1]1. Determining the fatigue characteristics of the TangoBlackPlus material. 2. Determining the fatigue characteristics of the interface between TangoBlackPlus and another stiffer material (VeroWhitePlus).3. Determine factors that may increase or compromise the fatigue life of the printed TangoBlackPlus material.4.

Determine factors that may increase or compromise the fatigue life of the printed TangoBlackPlus material.4. Offering design guidelines to improve the fatigue life of the printed elastomeric components.To accomplish these goals, the authors first provide a review of the literature in Section 2, describe the experimental methodology in Section 3, present and discuss the results of the experiments in Section 4, and finally describe the implications of the results and discuss areas for further research in Section 5.2. Literature Review and Theory2.1 Fatigue in Printed ElastomersThe manufacturer does not offer any information on the fatigue properties of the printed elastomeric material [1].

Literature Review and Theory2.1 Fatigue in Printed ElastomersThe manufacturer does not offer any information on the fatigue properties of the printed elastomeric material [1]. Though information is available on the static material properties of TangoBlackPlus [1] and work has been done to model the static properties of direct 3DP materials [5, 6], a literature search did not reveal any research examining the fatigue characteristics of direct 3DP components.Because no research thus far has examined the fatigue characteristics of printed elastomers, two bodies of literature were examined to better understand the theory applicable to fatigue failure in printed elastomers: (i) literature on the fatigue characteristics of traditionally manufactured elastomers and (ii) literature on the fatigue characteristics of non-elastomer AM642parts.

Together it is hoped that the two bodies of literature will provide some insight into the fatigue behaviors of printed elastomers.2.2 Fatigue in Traditionally Manufactured ElastomersBecause of the extensive use of rubbers in automobile tires, which experience high cycle loadings during regular use and pose a safety risk upon failure, the fatigue properties of many rubber compounds have been extensively examined. The tensile fatigue failure process in rubber as documented by the literature [3, 4], begins at small imperfections, usually located on the surface of the specimen that are introduced through the manufacturing process. These small imperfections cause local stress concentrations under loading. If stresses become large enough in these stress concentrations, some of the polymer chains can be mechanically broken, expanding the imperfection.

If stresses become large enough in these stress concentrations, some of the polymer chains can be mechanically broken, expanding the imperfection. Upon repeated loadings, the tear in the material is lengthened by a small amount each cycle until, at some point, the specimen experiences catastrophic tearing where the tearing rate increases substantially and complete part failure is the eventual result. Through this process, repeated cyclic exposure to loads significantly less than the maximum static load can result in complete part failure.Below a given strain, typically less than 65% - 95% elongation for vulcanized rubbers, the elastomers can exhibit a fatigue limit [7]. Below this elongation, stress concentrations are not large enough to mechanically rupture the polymer chains around the part imperfections. This leads to a sharp jump in fatigue life, so long as the given strain is not exceeded in the cyclic loading.

This leads to a sharp jump in fatigue life, so long as the given strain is not exceeded in the cyclic loading. Parts experiencing cyclic loading below this limit do eventually experience complete part failure, but the primary driver is the chemical degradation of the material, rather than mechanical rupture of the polymer chains.2.3 Fatigue in Non-Elastomeric AM PartsThere is some research on the fatigue properties of AM parts, but the literature is scarce. If a goal of the AM community is to use the technology as a manufacturing process, rather than just a prototyping tool, more research on the fatigue properties of AM components is needed. The existing research that was found on fatigue in AM components focuses on direct metal AM processes.Studies have shown that if parts are sufficiently dense, and they are oriented to prevent delamination, the fatigue life of both plastic [8] and metal AM components [9, 10] can match their traditionally manufactured analogs.

Different studies however have found that AM components have significantly shorter fatigue lives than their traditionally manufactured analogs [11, 12]. The presence of internal voids [9–12], and lamination weaknesses between build layers [8] in the AM components have been linked to premature fatigue failure of parts. Santos et al. inversely correlated the number of pores in metal AM specimens to the fatigue life of those specimens [12].6433. MethodsFollowing a review of the AM literature, it was found that no prior literature existed on the fatigue life of printed elastomers. To address this gap in the literature, the authors performed fatigue testing of multiple elastomeric specimens fabricated with the Objet Polyjet process.

To address this gap in the literature, the authors performed fatigue testing of multiple elastomeric specimens fabricated with the Objet Polyjet process. The methods used for this experiment are described in Section 3; the results of the experiment are presented in Section 4.3.1 Specimen Creation ProcedureBecause the printed elastomer is labeled as “rubber-like” by the manufacturer [1], the ASTM D4482-11 “Test Method for Rubber Property Extension Cycling Fatigue” [13] was used to guide the fatigue testing procedures for this study of the Objet TangoBlackPlus material.The fatigue testing specimens were designed according to the ASTM standard and consisted of a modified dogbone shape. The exact dimensions of this shape can be seen in Figure 3. The specimen takes the shape of the traditional dogbone specimen with the addition of beaded edge, designed to fit into a set of special elastomer grips.

The specimen takes the shape of the traditional dogbone specimen with the addition of beaded edge, designed to fit into a set of special elastomer grips. This feature is necessary with elastomers because the lateral shrinkage of the specimens under high strains tends to let specimens pull out of traditional grips.Figure 3: An engineering drawing of the fatigue test specimens. All measurements given in millimeters.All specimens were printed on an Objet Connex 350 and cleaned of support material in a water-jet station. Specimens were printed with the TangoBlackPlus material in “digital material” mode (i.e., 32 m layer thickness). The surface finish was set to the “matte finish” for all parts; thus, all surfaces were encased in support material. This ensured that the surface finish was consistent across all surface of the test specimens. The specimens were printed flat in the XY plane with the longest dimension of the part in the direction of the movement of the print carriage.

The specimens were printed flat in the XY plane with the longest dimension of the part in the direction of the movement of the print carriage. Samples were positioned onto the build tray so as to ensure that all samples were printed with a single pass of the print head (per layer). Samples were positioned onto the build tray so as to ensure that all samples were printed with a single pass of the print head (per layer). This was done to ensure that all parts received uniform exposure from the UV lamp, (and therefore had uniform curing), since the differing UV exposure levels due to part spacing and part orientation can cause significant differences in part strength [14, 15].644To test the fatigue characteristics of the interface between the TangoBlackPlus material and a standard non-elastomer material (Objet’s VeroWhitePlus, in this case) two alternate specimen designs were used: a flat interface specimen (Figure 4a) and a shear interface specimen (Figure 4b).

Each of these specimens had the same overall part dimensions, but the specimens were made from the two materials.Figure 4: The flat interface and shear interface specimensThe flat interface specimen was intended to test the performance of the elastomer/non- elastomer interface under pure tensile loading. The interface between the VeroWhitePlus material and the TangoBlackPlus material was a flat plane normal to the direction of extension located in the reduced central region of the dogbone, five millimeters from the edge of the reduced central section. The interface was purposefully not centered, giving more area to the TangoBlackPlus, so that strain measurements could more accurately be made on the TangoBlackPlus material.The shear interface specimen was intended to test the performance of the elastomer/non- elastomer interface under a combination of shear and tensile loading.

Again, the interface was located in the reduced central region of the dogbone with five millimeters from the center of the interface to the edge. The interface was rotated by 45o from vertical, so that the interface plane would line up with the maximum shear plane in the specimen under extension.3.2 Specimen Testing Procedure:All specimens were printed and cleaned one day prior to testing to mitigate the effect of partageing in this study. After cleaning, parts were allowed to dry overnight before testing.Specimens were mounted and tested in an MTS Tytron 250 Microforce Testing System with a 250N load cell (Figure 5). The testing apparatus was chosen for its sensitivity to small loads and its large extension range. Special grips were designed and built on an FDM machine (from P-400 ABS material) to hold the beaded edge of the test specimens. The grips were adjusted until the load cell read between 0.05 and 0.1N of tension force in the specimen.

The grips were adjusted until the load cell read between 0.05 and 0.1N of tension force in the specimen. This was set as the zero strain point for the specimen.Figure 5: The MTS Tytron 250 Microforce Testing System645Because of delicate and non-rigid nature of the elastomer fatigue specimens, an extensometer could not be used to measure strain in the specimens. Instead, two small marks were painted onto the center section of the test specimen, at least two millimeters from the neck region of the specimen on either side. The distance between the inside edge of these two marks was measured with a set of calipers. The extension was then adjusted until the marks measured the appropriate reading for the desired maximum strain. The extension measurement and the force at the desired extension were then recorded.

The extension was then adjusted until the marks measured the appropriate reading for the desired maximum strain. The extension measurement and the force at the desired extension were then recorded. Pilot testing revealed that the relationship between extension and measured strain was very reliable across specimens, so the extension/strain coordination measurement was conducted only once at each desired level of strain for each type of fatigue specimen.After the specimens were mounted and the maximum extension was determined for the desired strain, the specimens were tested in a loading and relaxation cycle until complete part failure. The loading and relaxation cycle was taken from the ASTM standard and consists of four phases. The standard specifies a testing frequency of 1.7 Hz, with four phases of equal length (each 0.147 sec). The phases are:1. phase one consisted of hold at zero extension. 2.

The standard specifies a testing frequency of 1.7 Hz, with four phases of equal length (each 0.147 sec). The phases are:1. phase one consisted of hold at zero extension. 2. Phase two consisted of a linear ramp from zero extension up to the maximum extension.3. Phase three consisted of a hold at the maximum extension. 4. A ramp down from the maximum extension to zero extension.Figure 6 shows the commanded extension path and the actual extension part for one specimen during testing.Figure 6: A plot of the fatigue extension cycle. In this example maximum extension is set to 8mm.Complete part failure was defined as a complete rupture of the part, leaving the specimen in two pieces. Part failure was identified through an automated check of the tensile force at the beginning of phase three of the loading and relaxation cycle. If the tensile load during maximum extension fell below 0.1 N, the specimen was determined to have failed and the test was halted.

If the tensile load during maximum extension fell below 0.1 N, the specimen was determined to have failed and the test was halted. The number of cycles was then rounded to the nearest integer and the cycles to failure for the given specimen was recorded.6464. Results and Discussion4.1 Comparing the Fatigue Life of the Material and Multi-Material InterfaceBefore developing the ε-N curve, which relates the maximum strain (ε) to the expected material life in number of cycles (N), it was necessary to determine if the multi-material interface was weaker than the TangoBlackPlus material. Because the majority of example applications of TangoBlackPlus incorporate other materials interfacing with the TangoBlackPlus material [1], the fatigue characteristics of the multi-material interface were also of interest. It was hypothesized by the authors that the multi-material interface would fail under fatigue loading before the material itself failed.

It was hypothesized by the authors that the multi-material interface would fail under fatigue loading before the material itself failed. If this was true, the fatigue characteristics of the interface would be more valuable to the designer than the fatigue characteristics of the TangoBlackPlus material itself. To test this hypothesis, a comparison was carried out between test specimens made of only TangoBlackPlus, with no interface, and the flat interface specimens that did have a material interface.To determine if the multi-material interface was in fact weaker than the TangoBlackPlus material, a set of twelve non-interface specimens was compared to a set of twelve flat interface specimens and statistical differences were examined in the fatigue life of those specimens. Extensions were set so that the TangoBlackPlus material located in the reduced central section of dogbone specimens experienced 40% elongation (a mid-range extension value identified in pilot testing).

The results of this comparison are shown in Table 1.Table 1: A comparison of specimens with and without a multi-material interfaceNo InterfaceFlat InterfaceSpecimen Number 1 2 3 4 5 6 7 8 9 10 11 12Cycles to Failure568 1259 1003 787 1158 1226 807 1105 841 1020 1617 1367Break Locationneck neck neck neck neck neck neck neck neck neck neck neckCycles to Failure1794 1670 1090 1122 1299 1184 1063 1376 1388 1740 1289 583Break Locationneck neck neck neck interface neck neck interface neck neck neck materialMean Cycles to FailurePercentage Break at NeckMean Cycles to FailurePercentage Break at Neck1063100%130075%Standard DeviationStandard Deviation289337P-value* = .078 *Two tailed Students T-Test, no assumption of equal variance647Contrary to the hypothesis, it was found that the interface specimens actually had a higher fatigue life than the non-interface specimens (though this difference was not statistically significant at the p = 0.05 level).

Additionally, it was found that 100% of the non-interface specimens and 75% of the interface specimens failed at the necks made of TangoBlackPlus, rather than at the multi-material interface. Images of some of these specimens can be seen in Figure 7.This finding led the authors to form a revised hypothesis that the TangoBlackPlus neck in the specimen was the weak point of the design. To test this new hypothesis, a new set of specimens with two multi-material interfaces was made. A drawing of these new specimens can be seen in Figure 8.

To test this new hypothesis, a new set of specimens with two multi-material interfaces was made. A drawing of these new specimens can be seen in Figure 8. With no necks made of the TangoBlackPlus material, a comparison of this new set of specimens to the existing two sets of specimens would allow the authors to compare specimens with and without necks made of the TangoBlackPlus material.Figure 7: Pictures of the breaks at the TangoBlackPlus specimen neckFigure 8: The Dual Interface SpecimenThe extension for these new dual interface specimens was set so that the TangoBlackPlus materials experienced 40% elongation. Twelve specimens were tested at the 40% elongation level.

Twelve specimens were tested at the 40% elongation level. These results are outlined in Table 2.648Table 2: The Dual Interface SpecimensDual InterfaceCycles to Failure 32 1345 1209 2128 1226 3491 2588 2083 139 1068 372 250Break Location interface material material material material interface material material interface material interface interfaceComparison of dual interface and nointerface specimensp-value* = .426Comparison of dual interface and nointerface specimensp-value* = .933Two tailed Students T-Test, no assumption of equal varianceMean Cycles to FailureStandard Deviation13281076As is seen in Table 2, the dual interface specimens had higher fatigue lives on average than both those with no interface and the single interface specimens (Table 1); however, neither of the differences are statistically significant. It can also be observed that the dual interface specimens seemed to have more variation in terms of the fatigue life.

It can also be observed that the dual interface specimens seemed to have more variation in terms of the fatigue life. Finally, it was observed that 58% of the dual interface specimens failed in the central region (TangoBlackPlus), and not at either interface. Based on this data, there is no evidence to support the hypothesis that the multi- material interface has a fatigue life that is inferior to the TangoBlackPlus material’s fatigue life.4.2 Testing the Effects of Shear on the InterfaceThe multi-material interface was not found to have a shorter fatigue life than the TangoBlackPlus material under loading normal to the interface plane; however, this result was for a pure tensile loading of the interface. In application, the interface may be subjected to shear loading as well. To determine the effect of shear loading on the fatigue life of the specimen, another specimen was developed.

In application, the interface may be subjected to shear loading as well. To determine the effect of shear loading on the fatigue life of the specimen, another specimen was developed. This specimen was identical in design to the dual interface specimen, except one of the interfaces was angled 45o from vertical. This lined the interface up plane under the maximum shear force when the specimen was under elongation.Twelve of these new shear specimens were created and compared to the original dual interface specimens. The extension for these new shear interface specimens was set so that the TangoBlackPlus materials experienced 40% elongation measured along the longer side of the TangoBlackPlus material.

The extension for these new shear interface specimens was set so that the TangoBlackPlus materials experienced 40% elongation measured along the longer side of the TangoBlackPlus material. The results of this experiment are outlined in Table 3.649Table 3:Shear interface specimensShear InterfaceCycles to Failure 474 500 490 585 209 733 1221 692 1100 776 855 488Break Location shear interface shear interface shear interface shear interface shear interface shear interface shear interface shear interface material shear interface shear interface shear interfaceComparison of flat interface specimens to the shear interface specimens p- value* = .065Two tailed Students T-Test, no assumption of equal varianceMean Cycles to Failure677Percentage Break at Shear Interface 92%Standard Deviation285As is seen in Table 3, the introduction of the angled interface nearly halves the average fatigue life of the specimens.

This difference is not statistically significant at the p = 0.05 level because of the large spread of the data, which is common in fatigue life analysis [16]. It can be observed that all but one of the shear specimens broke at the angled interface, indicating that the angled interface may indeed be the weakest point of the specimen.4.3 Relating Material Strain to Expected Fatigue LifeUsing the normal dual-interface specimens (Figure 8), which were identified as the most representative of the TangoBlackPlus material properties, the effect of maximum elongation on fatigue life was examined next. To relate these two variables, the authors sought to create a ε-N curve (also known as the Wӧhler curve), to correlate material strain to expected fatigue life.Twelve tests were conducted at each of the desired elongation ratios (20%**, 30%, 40%, 50% and 60%), to build this curve.

The minimum value of 20% elongation was chosen because specimens were observed in pilot tests to have achieved fatigue lives of more than one million cycles at this extension ratio, which was an upper limit set on test duration set by the authors. The maximum of 60% elongation was chosen because specimens at 70% elongation regularly failed on the first elongation (indicating a maximum static elongation at break.)A scatterplot of the data points can be seen in Figure 9. The break location of samples isindicated by the color of the point.650% Elongation vs. Fatigue Life7060InterfaceMaterial50Did Not Breakn o i t a g n o l E %403020100101001000100001000001000000Cycles to FailureFigure 9: A scatterplot showing the relationship between material strain and fatigue life. Observed failure point indicated by color.The scatterplot provides insights into the general shape of the ε-N curve, as well as the nature of the interface and the material.

Observed failure point indicated by color.The scatterplot provides insights into the general shape of the ε-N curve, as well as the nature of the interface and the material. As can be seen in the data points, the specimens that failed in the material seem to have fatigue lives that are clustered fairly tightly around a central value. Alternatively, the fatigue lives of the specimens that broke at the interface are more erratic. The interface sometimes failed very prematurely, other times it did not fail at all (the material itself failed instead.)A curve that predicts the fatigue life of a specimen for a given strain was constructed by averaging the scatter plot data of Figure 9, and is presented in Figure 10. The 95% confidence interval is denoted by the dashed lines in Figure 10.6516050n o i t a g n o l E %403020101001000100001000001000000Expected Fatigue LifeFigure 10: The ε-N curve for TangoBlackPlus.

95% CI denoted by the dashed lines.Figure 10 gives designers a prediction on the fatigue life of their printed TangoBlackPlus component, based on the expected strain the material will experience. The confidence interval covers a very broad range, but this wide variability is a typical characteristic [16] of fatigue. Additionally, the erratic nature of the confidence interval is noteworthy. This is primarily due to the few very premature failures exhibited in the data (Figure 9). Elongation ratios that had more premature failures (such as 40% elongation) have wider variability, while elongation ratios that had only one or two premature failures (such as 30% elongation) have much less variability. With larger sample sizes, the erratic confidence interval would most likely be smoothed out.5. Closure and Future Work5.1 ConclusionsIn this paper, the authors characterized the fatigue properties of a 3D printed elastomer(specifically, Objet’s TangoBlackPlus).

Closure and Future Work5.1 ConclusionsIn this paper, the authors characterized the fatigue properties of a 3D printed elastomer(specifically, Objet’s TangoBlackPlus). The findings are as follows:Contrary to the authors’ hypothesis, the multi-material interface between elastomer and non-elastomer did not appear to have a shorter fatigue life than the material itself on average. This being said, there was some indication that the interface had more erratic fatigue behavior than the material, occasionally causing very premature fatigue failure.  Reductions in in cross sectional area, such as the neck of the fatigue specimens, may cause serious reduction in fatigue life, even if the reductions are gradual.652Shear forces on the multi-material interface have a negative impact on the fatigue life of the multi-material interface.A relationship between strain and fatigue life was found (Figure 10).

Despite voids that are inherent to the 3DP process, the material specimens exhibited a long life at low extension ration (up to 106 cycles at 20% elongation)5.2 Design Advice for Improving the Fatigue Life of Printed ElastomersBased on the above conclusions, the authors offer the following suggestions for designersusing 3D printed elastomers, in order to maximize the life of the elastomers their product.1. Reduce the strain exerted on the elastomer. It was found that under low strains (20% elongation or less) the elastomers had a long fatigue life, approaching 106 cycles.2. Avoid putting the elastomer/non-elastomer interface under shear loads. These shear loads will shorten the fatigue life of the multi-material interface.3. Do not add necks to increase the area of the multi-material interface.

These shear loads will shorten the fatigue life of the multi-material interface.3. Do not add necks to increase the area of the multi-material interface. This will not increase the design’s life; it may in fact shorten the lifetime of the product.5.3 Future WorkWhile this work provides insight into the fatigue characteristics of printed elastomers (e.g., fatigue life, material interface, interface geometry), there still remains opportunity for further research:What are the compressive fatigue characteristics of printed elastomers? This study only examined the tensile fatigue properties, a study that investigates compressive fatigue characteristics is needed.What is the effect of build orientation of the fatigue characteristics of printed elastomers? Only one build orientation (X-Y plane) was used in this study to simplify the analysis.

 What are the effects of surface finish on the fatigue life of printed elastomers? Different surface finish options may improve or harm the fatigue life of the material.  What is the effect of part ageing the fatigue printed characteristics elastomers? Past research has shown that photopolymers degrade with age [5].ofWhat is the effect of strain rate on the fatigue life of printed elastomers? Figure 11 shows an observed hysteresis effect with the fatigue specimens. This indicated a non-linearity in the stiffness of the material.Figure 11: A plot showing the hysteresis effect in the TangoBlackPlus material6536. AcknowledgementsThe authors would like to thank Nicholas Meisel for his help printing specimens, Dr. Dan Dudek for the use of his laboratory and the MTS Tytron 250, and Albert Kwansa for his expertise and help in using the MTS Tytron 250.7.

TIGUE CHARACTERIZATION OF ULTRASONIC ADDITIVE MANUFACTURED ALUMINUM 3003P.J. Wolcott and M.J. DapinoDepartment of Mechanical and Aerospace Engineering, The Ohio State University, Columbus, OH, 43210AbstractAn aluminum 3003 H-18 block was built using ultrasonic additive manufacturing with process parameters which optimize mechanical strength. Transverse tensile fatigue tests were conducted on samples cut from the block and a stress vs. number of cycles curve was generated. Results show the curve is relatively flat and a stress threshold of 50% of the ultimate transverse tensile strength exists below which failure does not occur within 3.75 × 107 cycles. Optical and scanning electron microscopy conducted on failure surfaces shows no signs of crack initiation or growth typical of fatigue loading, but exhibits areas of no bonding between foil layers. To explain the failure, a model was developed based on the probabilistic failure of bonded areas in tension.

To explain the failure, a model was developed based on the probabilistic failure of bonded areas in tension. The model uses a Frechet distribution to describe the probability of failure at each individual bonded area. Discrete bond failure and subsequent redistribution of stress to the surviving areas eventually leads to failure for the entire sample. Predictions from the model show good correlation with the experimental results.IntroductionUltrasonic additive manufacturing (UAM), also known as ultrasonic consolidation, is asolid state manufacturing process that joins layered metal foils using principles of ultrasonic metal welding in combination with subtractive milling processes to create near net shape metal parts [1]. The process uses a rolling sonotrode to apply ultrasonic vibrations to metal foils; the scrubbing action produced by the sonotrode breaks up oxides and contaminants on the surface of the adjacent foil layers generating nascent bond surfaces.

Normal force applied via the sonotrode creates intimate metal to metal contact between the nascent surfaces which results in metallic bonding well below the foil melting temperature [2]. Heat generation due to the scrubbing action ranges from 30% to 50% of the foil melting temperature [3]. Therefore, UAM provides unprecedented opportunities to manufacture metal-matrix composites with embedded elements such as smart materials and electronics that are normally sensitive to thermal cycling [4] [5]. Input parameters of tack force, weld force, weld rate, amplitude and preheat temperature can be adjusted depending on the foil material to maximize the bonding and mechanical strength between foil layers [6].To realize the entire potential of UAM, it is necessary to fully characterize the mechanical properties of UAM builds.

Several studies have utilized microstructure, peel strength or void density analyses to understand the relationship between bond quality and manufacturing parameters, focusing mainly on aluminum 3003 as the foil material [7] [8] [9]. These tests provide valuable information on the bonding characteristics of UAM builds, but are difficult to correlate to common mechanical strength properties. Studies by Hopkins et al. [6] [10] have428determined optimal sets of parameters for various material combinations based on mechanical strength testing, including transverse tensile and shear strength. These tests provide a direct comparison to the mechanical properties of traditionally manufactured metal structures. Specifically, transverse tensile tests provide an indication of the strength at the interface with the lowest bond strength.Fatigue characterization of UAM built structures is important to determine potential lifetimes in cyclic loading conditions for structures in service.

Up to this point, fatigue analyses of UAM built structures have not been available in the literature. The fatigue testing presented here will provide the first such lifetime prediction data for the design of UAM components.Experimental ProcedureSample FabricationA UAM block with dimensions 101.6 mm x 50.8 mm x 33 mm (4 in. x 2 in. x 1.3 in.) was fabricated using a Solidica, Inc. Beta UAM machine (2 kW) at Edison Welding Institute, Columbus, Ohio from which fatigue samples were cut. The block consisted of Al 3003 H-18 foils, 152.4 μm thick (0.006 in.) built upon a 12.7 mm (0.5 in.) thick Al 3003 H-14 baseplate. Joining of successive foil layers was accomplished through tacking and welding passes; the input energy in the welding pass was higher than in the tacking pass. A staggered, alternating layup of foils was used to eliminate seams running through the height of the build.

A staggered, alternating layup of foils was used to eliminate seams running through the height of the build. In addition, milling passes were conducted periodically throughout the building process to maintain flatness. Samples were cut from the block via wire electronic discharge machining avoiding any seams and machined via CNC lathe to the specified dimensions shown in Figure 1. Sample dimensions include 12.7 mm baseplate material to more closely follow dimensional requirements in ASTM standard E466-7 for conducting axial loaded fatigue tests [11].The process parameters used for the build are given in Table 1. The parameters chosen were based on a study by Hopkins et al. [10] that determined optimal process parameters for mechanical strength of UAM Al 3003 H-18 builds on the Solidica Inc. Beta UAM machine.

The parameters chosen were based on a study by Hopkins et al. [10] that determined optimal process parameters for mechanical strength of UAM Al 3003 H-18 builds on the Solidica Inc. Beta UAM machine. These parameters provide an ultimate shear strength of approximately 22 MPa and an ultimate transverse tensile strength (UTTS) of approximately 30 MPa when using Al 3003 H-18 foils.Figure 1 - Sample dimensions for transverse axial fatigue testing (foils not to scale).429Table 1 - Process parameters for UAM build.Tack Force Weld Force Amplitude Weld Rate Tack Rate Tack Amplitude Base Plate Temperature Sonotrode Texture, Ra Oscillation Frequency350 N 1000 N 26 μm 42 mm/s 50.8 mm/s 9 μm 150°C 7 μm 20 kHzSample TestingFatigue testing was conducted using principles outlined in ASTM standard E466-7 [11]in a MTS 831 test frame with 25 kN load capacity specially designed for high-cycle fatigue testing.

Samples were loaded axially in tension with a minimum load of 22.24 N (5 lbs) to avoid compressive strokes that could affect fracture surfaces after failure. The R-ratio, given by,(1)was near zero for these tests. Loading was conducted transverse to the foil layup direction as seen in Figure 1. Since the transverse orientation is the lowest strength direction for UAM composites [12], it is necessary to study the fatigue behavior in this orientation. The cyclic tests were conducted using load control with the number of cycles counted until failure. The frequency of cycling for each test was ramped to 10 Hz within 100 cycles and sustained through 1.0 × 106 cycles. For samples tested beyond 1.0 × 106 cycles, the frequency was increased to 30 Hz to expedite the testing. This test procedure and the frequencies used are well within the standards and are commonly used for fatigue load cycling [11] [13].

This test procedure and the frequencies used are well within the standards and are commonly used for fatigue load cycling [11] [13]. After testing, all of the samples were examined via optical and scanning electron microscopy (SEM) to investigate the specimen failure surfaces for bonding and failure characteristics.Experimental ResultsThe stress-number of cycles (S-N) curve showing the results of the fatigue testing isgiven in Figure 2. The plot shows the maximum stress vs. cycles to failure as well as the percentage of the UAM ultimate transverse tensile strength for each stress level. Samples were considered runouts when loaded to 3.75 × 107 cycles without failure; the three samples that reached this threshold are designated with arrows in Figure 2.

Samples were considered runouts when loaded to 3.75 × 107 cycles without failure; the three samples that reached this threshold are designated with arrows in Figure 2. Two of these samples were tested again at higher stress levels to maximize the number of tests in the study; these retests are represented with R on the figure.Two samples which failed in unexpected ways were omitted from the results. In Figure 2, these outliers are indicated with circles around the data points. One of the samples failed along a mill line that was created due to the flattening passes during the build, seen in Figure 3(a). This430failure is attributed to a manufacturing defect caused due to mill pass creating a height difference along a build interface. Another sample failed near the first layer at the edge of the gauge region, seen in Figure 3(b), indicating that the first layer did not weld strongly to the baseplate.

Another sample failed near the first layer at the edge of the gauge region, seen in Figure 3(b), indicating that the first layer did not weld strongly to the baseplate. Since these failures arise from identifiable and atypical manufacturing issues, these two results are disregarded in the fatigue analysis.Fatigue fracture surfaces for the samples that failed by delamination were similar to those found in [10] for UTTS samples; an example optical microscope image is shown in Figure 4. Images from the top foil layer of the delamination failure were utilized because they are not textured directly by the sonotrode during the building process, providing an ideal surface for bond characterization. The failure surface shows dark areas as original foil texture and lighter areas that have been textured by the foil layer it was bonded to. Important features not present in the optical microscopy are a clear crack initiation site and striations growing from this site.

Important features not present in the optical microscopy are a clear crack initiation site and striations growing from this site. This indicates a need to look more closely at the failure surfaces using scanning electron microscopy.2580) a P M(s s e r t S x a M2015105R-AR-BB A70605040302010001E+01E+11E+21E+31E+41E+51E+61E+71E+8Cycles (N)Figure 2 - Stress-cycle plot of maximum stress and cycles to failure. The arrows represent specimens that did not fail within 3.75 × 107 cycles and R represents samples retested at higher stress levels.

The arrows represent specimens that did not fail within 3.75 × 107 cycles and R represents samples retested at higher stress levels. Circled points represent samples that did not fail by delamination of foil layers.431S T T U M A U %(a)(b)Figure 3 - (a) Optical microscope image of sample failure along milling interface with mill marks shown; (b) Sample failure at edge of gauge region at the baseplate layer.Example SEM micrographs in Figure 5 show areas of ductile failure typical of metallic bonding, areas of original foil texture indicating no bonding, and areas of foil to foil contact without bonding (designated as flow), each consistent with the optical and SEM results presented in [10]. These characteristics were present in all of the failure surfaces. The areas of original foil texture are seen in the images as flat and dark whereas the ductile areas are lighter in color with a cup- cone like structure.

The areas of original foil texture are seen in the images as flat and dark whereas the ductile areas are lighter in color with a cup- cone like structure. The areas of foil to foil contact without bonding are lightly colored as well but do not show the cup-cone structure. None of the optical or SEM microscopy analyses show the characteristics expected in a fatigue failure, such as a well-defined crack initiation site and fatigue striations showing the growth of an initial crack.Figure 4 - Optical micrograph of fatigue fracture surface.432(a)(b)Figure 5 - SEM images of fatigue fracture surface showing ductility, foil to foil contact without bonding (flow), and original foil texture: (a) 312x and (b) 625x.Discussion of Experimental ResultsThe S-N curve suggests that a threshold of approximately 50% of the UAM ultimatetransverse tensile strength exists below which fatigue failure does not occur within 3.75 × 107 cycles.

Because it is known that nonferrous metals such as aluminum do not have a fatigue limit, it is expected that UAM Al 3003 builds would fail eventually given sufficient cyclic loading. At 30 Hz, testing a sample to 3.75 × 107 cycles requires more than two weeks. For this reason, more stress cycling on the samples that ran out in this study was not conducted. The S-N curve for the UAM Al 3003 build is relatively flat with most samples failing in the range from 20 to 15 MPa. These results, in addition to the lack of classic fatigue characteristics indicate a need for alternate explanations for lifetime prediction of UAM samples. It is therefore necessary to investigate alternate explanations for the failure using scaled modeling efforts.Model DevelopmentThe absence of classic fatigue features, such as a crack initiation site or striations, motivates the development of an alternate theory to describe the failure of UAM samples under dynamic loading conditions.

The model developed in this case uses the concept of simultaneous tensile loads on the discrete bonded areas of a UAM build. As each loading cycle is applied, there is a probability of failure for each bond that, if failure conditions are met, can lead to failure of that bond and a redistribution of stress to the surviving bonded areas. Using a probability of failure approach incorporates the concepts of stress concentrations, stress buildup, and asymmetric loading conditions under one metric. Simulating the failure in this way avoids having to calculate each phenomenon individually with a complex stress analysis at each bond. This makes the model ideal as an initial test without being overburdened with unknown factors.

This makes the model ideal as an initial test without being overburdened with unknown factors. Figure 6 shows a side view schematic of the bonding illustrating the model theory, whereby on each load cycle there is a possibility of bond failure.433The evolution of stress on the bonds is described by,(2)where F0 is the nominal load applied, A0 is the nominal area of the sample (not accounting for actual bonding), is the percentage of actual bonding present prior to loading (18.1% [10]), and x is the percentage of surviving bonds starting at an initial value of 100% without any load cycling and 0% at failure. This equation therefore scales the stress based on the amount of bonding present during the loading. The model uses a set of random numbers checked against the probability of failure curve to determine whether a bond is to fail or not. In this case a random number from 0-1 is checked against the probability curve at the loading stress.

In this case a random number from 0-1 is checked against the probability curve at the loading stress. If the random number is less than the probability of failure at a given stress level, the bond is deemed to have failed and its value becomes zero.For simplicity, the bonding was assumed to be 100 equally sized areas. The value for x is the sum of a 100 index vector, each initialized with a value of 1/100. As the bonds fail, corresponding indices are reset to 0, changing the sum of the vector to a value less than 1. This in turn decreases the number of surviving bonds, x, and therefore the stress increases according to equation (2). This process is continued until a condition is met where σ ≥ σUTS, where σUTS is the ultimate tensile strength of the wrought Al 3003-H18 alloy (200 MPa) [14]. When failure occurs, the number of cycles is reported as the final value for the fatigue life at the given initial stress.

When failure occurs, the number of cycles is reported as the final value for the fatigue life at the given initial stress. The use of random numbers and the range of probabilities of failure build considerable variation into the model to simulate the stochastic nature of UAM bonding and failure.The probability of failure curves were fit from the experimental results based on a Frechet distribution given by,( ) [ ()] (3)where s is a scale parameter, σ0 is a location parameter, and α is a shape parameter [15]. The model parameters for the failure probability are given in Table 2 and the curves used for the model are given in Figure 7. Two curves were generated using two values of σ0 to simulate the stochastic nature of the bonding in UAM structures.

Two curves were generated using two values of σ0 to simulate the stochastic nature of the bonding in UAM structures. In the model, the value of the probability of failure is selected as a random value between the two curves.FoilBondsBroken bondFoil(a)(b)Figure 6 - Schematic of tensile loading and bond fracture for model theory showing a side view of the bonded foil layers at: (a) Initial bond conditions and (b) Bonding after a load cycle.434Table 2 - Model parameters for Frechet distribution.Model Parameter s σ0 αValue 3 0.65, 0.80 5.2500150500.62000.40.811000.2Stress (MPa)Prob. of FailureFigure 7 - Frechet curve [15] of failure probability for bonded areas.This concept, and likewise the distribution, is similar to more commonly used techniques where a Weibull distribution is used to describe the probability of failure for a given material [16].

In this case however, each individual bond is investigated, instead of the failure of the part as a whole. For this model, the Frechet distribution was used because it provides an adequate description of the experimental fatigue data.Model ResultsThe results of the model are shown with the experimental results in Figure 8. Ten simulations were performed at each of the experimentally tested stress values. The two samples that failed atypically, as described previously and found in Figure 2, were omitted as outliers. It is emphasized that the stress values reported are recalculated based on equation (2), accounting for the bonded area percentage as bearing the load. The results of the model show good correlation with the experimental results. Notably, the model accurately predicts the sample runout for the three sets of experiments run at less than 80 MPa (adjusted stress).

Notably, the model accurately predicts the sample runout for the three sets of experiments run at less than 80 MPa (adjusted stress). These results indicate that the concept of probabilistic failure of bonded areas is a potentially suitable description of the behavior of low power UAM samples under dynamic loading conditions.435120) a P M10080110 MPa100 MPa(s s e r t S604091 MPa81 MPa2071 MPa0 1E+01E+21E+4 Cycles1E+61E+862 MPaExperimentalFigure 8 - Comparison of experimental and model results. Arrows indicate samples that did not fail within 3.75 × 107 cycles.Future WorkFurther refinements will be necessary to develop the model to more accurately describethe fatigue loading conditions by adjusting the parameters in the failure distribution. This will be accomplished through constraint optimization techniques to most accurately fit the model parameters to the known experimental values.

This will be accomplished through constraint optimization techniques to most accurately fit the model parameters to the known experimental values. Further experimental testing will be necessary as well to validate the accuracy. This will ideally come from additional UAM builds using Al 3003 and other materials. The current model is developed from the experimental data gleaned from this study. By creating additional builds using other materials, a wide scale lifetime prediction model can be developed for fatigue loading of UAM structures.Additionally, based on the concepts and initial model developed, a macro-scale model can be developed for static tensile loading conditions.Also, using these concepts, a set of fatigue experiments will be necessary for lifetime prediction in very high power (VHP) UAM.

Based on the results of the modeling efforts shown above, this type of analysis could be applied to VHP UAM structures and compared with the experimental results.ConclusionsUltrasonic additive manufacturing was used to build Al 3003 H-18 samples for fatiguecharacterization. Samples were tested in the transverse tensile geometry cyclically without compressive cycling. A relatively flat S-N curve was generated. Two samples were omitted from the fatigue results after attributing their failure to specific manufacturing issues. It may436therefore be necessary to investigate the frequency with which these defects occur and determine how to limit their occurrence during processing.In addition, a stress threshold of 50% of the UTTS was discovered below which failure did not occur within 3.75 × 107 cycles.

This provides a lifetime prediction where, if these conditions are not exceeded, it is expected that a UAM component built with these process parameters would not fail.Due to the lack of classic fatigue features in the failed specimens, a model was introduced to describe the failure as a simultaneous tensile loading of discretely bonded areas. The failure was modeled as a probability of failure using a Frechet distribution, whereby as the bonds in the UAM sample fail, the stress is redistributed and increased among the remaining bonds. The model shows a good correlation with the experimental results, providing an initial starting point for more refined methods to be applied to UAM bonded structures.AcknowledgmentsThe authors would like to thank Dr. S. Suresh Babu of the Welding Engineering Programin the Department of Materials Science and Engineering at The Ohio State University for his assistance.

S. Suresh Babu of the Welding Engineering Programin the Department of Materials Science and Engineering at The Ohio State University for his assistance. Financial support for this research was provided by the member organizations of the Smart Vehicle Concepts Center (www.SmartVehicleCenter.org), a National Science Foundation Industry/University Cooperative Research Center..References[1] K. Graff, "New Developments in Advanced Welding," Woodhead Publishing Limited,2005.[2] D. White, "Ultrasonic consolidation of aluminum tooling," Advanced materials &processes, vol. 161, no. 1, pp. 64-65, 2003.[3] C. Kong, R. Soar and P. Dickens, "Optimum process parameters for ultrasonic consolidation of 3003 aluminum," Journal of Materials Processing Technology, vol. 146, no. 2, pp. 181- 187, 2004.[4] E.

1-12, 2010.[8] T. Friel, K. Johnson, P. Dickens and R. Harris, "The effect of interface topography forultrasonic consolidation of aluminum," Materials Science and Engineering A, vol. 527, pp.4374474-4483, 2010.[9] M. Kulakov and H. Rack, "Control of 3003-H18 Aluminum Ultrasonic Consolidation,"Journal of Engineering Materials and Technology, vol. 131, 2009.[10] C. H. Hopkins, P. J. Wolcott, M. J. Dapino, A. G. Truog, S. S. Babu and S. A. Fernandez,"Optimizing ultrasonic additive manufactured Al 3003 properties with statistical modeling," Journal of Engineering Materials and Technology, vol. 134, no. 1, 2012.[11] ASTM E 466-07: Standard Practice for Conducting Force Controlled Constant AmplitudeAxial Fatigue Tests of Metallic Materials, ASTM International, 2007.[12] D. Schick, R. Hahnlen, R. Dehoff, P. Collins, S. Babu, M. J. Dapino and J.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFatigue Endurance Investigation of Post-processed Surfaces of L- PBF Ti-6Al-4V under Flexural StressC. Banuelos1,2*, B. Ramirez1,2, A. De la Cruz1,2, S. T. Nabil1,2, E. G. Arrieta1,2, R. B. Wicker1,2, and F. Medina1,21 Department of Aerospace and Mechanical Engineering, University of Texas at El Paso, El Paso, TX 79968, USA;2 W.M. Keck Center for 3D Innovation, El Paso, TX, USA;Corresponding author (cbanuelos4@miners.utep.edu)AbstractNumerous research works can be found focusing on fatigue properties of AM components, however most of this literature is focused on uniaxial testing. Because the very few actual components under uniaxial loading conditions found in any application, it is also important to investigate fatigue performance under loads that produce combined stresses, such as bending.

This project investigates the fatigue endurance of LPBF Ti-6Al-4V specimens subjected to four different surface finishing prost-processes (milled, ground, polished and abrasive media). The test consisted of a force-controlled cyclic load applied on the specimen in a 4-point bending setup until fracture. The study incorporated mechanical and optical techniques to measure and quantify the characteristic surface roughness of the post-processes. Additionally, failure mechanisms are discussed on fractographs. The data analyses suggested that internal defects commonly present in additively manufactured parts had a more significant impact on the fatigue life than surface roughness of post-processed parts.IntroductionAdditive manufacturing, commonly known as 3D printing, is a groundbreaking technology that enables the fabrication of intricate parts and objects by layer-by-layer construction of materials.

Laser Powder Bed Fusion (LPBF) is one of the most popular and versatile additive manufacturing techniques. It involves utilizing a high-powered laser to selectively melt and fuse metal or plastic powder, resulting in the desired shape formation. LPBF finds extensive applications in industries such as aerospace, automotive, medical, and consumer goods due to its ability to produce highly precise and customized parts [1-3].The focus of interest in this study lies in the manufacturing of aerospace-grade titanium alloy (Ti-6Al-4V) using LPBF. This material possesses numerous advantages, including lightweight properties and high strength. However, the LPBF process introduces certain drawbacks, such as residual stresses and poor surface finish. These factors can adversely affect the mechanical performance and fatigue life of the manufactured parts [4,5].

These factors can adversely affect the mechanical performance and fatigue life of the manufactured parts [4,5]. Various post-processing techniques, including heat treatments and machining methods, can be employed to mitigate these1909disadvantages. However, it is important to note that these processes often come with added costs and delays in the supply chain.Previous research studies [6-8] have demonstrated a correlation between surface quality and fatigue performance under cyclic loading. An inadequate surface finish creates imperfections that lead to stress concentration, ultimately causing the part to fail. In the context of this research, poor surface finish can serve as initiation sites for crack growth, leading to reduced fatigue performance.The objective of this study is to evaluate the impact of surface finish on the performance of additively manufactured parts and quantify the extent to which bending fatigue life can be enhanced.

Additionally, the study aims to determine if specific post-processing techniques offer significant advantages over others in terms of improving fatigue performance.MethodsPowder characterizationThe specimens used in this research were manufactured using grade 5 titanium alloy powder. The powder was sampled according to ASTM B215-15 standard [9] and was then characterized by a Retsch Technology Camsizer X2 X-Dry as per ASTM B822 standard [10].Figure 1: SEM image of Ti-6Al-4V powder1910Figure 2: Powder size distribution graphEOS build parametersThe specimens were built in an EOS M290 laser powder bed fusion system equipped withtwo Ytterbium lasers in an Argon environment. All samples were built using EOS Nominal parameters; that is a laser power of 280 W at a scanning speed of 1200 mm/s.

All samples were built using EOS Nominal parameters; that is a laser power of 280 W at a scanning speed of 1200 mm/s. These parameters are located within the process window as demonstrated in Figure 4 below.Figure 3: Image of EP08 samples on build plate1911Figure 4: P-V diagram of the process regions of grade 5 titanium for L-PBFHeat treatmentThe parts were stress relieved to remove any internal thermal stress commonly associated with the laser powder bed process. The stress relief was achieved by allowing the parts to evenly heat up to 600°C and then evenly cool down to room temperature in a vacuum furnace at a ramp and cooling rate of 5°C/min.Sampling and test developmentThe parts were then randomly assigned into the different machining strategies as well as designated toward developing the test design.

Furthermore, the machined samples were allocated into four different maximum stress levels (667 MPa, 900 MPa, 1067 MPa, and 1200 MPa) for testing and three samples from each machining technique were set aside as spares as shown in Table 1. Figure 5 illustrates the print location of the samples and how they were randomly assigned to the different surface finishes.Table 1: Test MatrixMachining MethodTangential Milling Deci Duo Surface Ground LTI Polished Total1200 MPa 3 3 3 3 121067 MPa 3 3 3 3 12900 MPa 3 3 3 3 12667 MPa 3 3 3 3 12Spares3 3 3 3 12Total15 15 15 15 601912Figure 5: Map of sample assignments according to their build plate print location.Machining techniquesFour different machining techniques were explored in this project, the first of which was a thermal atomized fusillade performed on a Deci Duo machine by Post Process.

The intent of this process is to reduce the processing time of AM parts by streamlining the steps from finished print to finished product. This method works by allowing a combination of detergent, abrasive media, water, and air to be shot out of nozzle as the sample rotates to improve the surface quality. Figure 6 illustrates how the surface of these samples looked once completed. Figure 7 demonstrates how the specimens were setup and the parameters used for this process.Figure 6: Image of finished Deci Duo sampleFigure 7:Illustration of Deci Duo processThe second machining technique was tangential milling. This process consisted of utilizing a half inch cutter to remove 0.5-mm of material from each side in a single pass.

This process consisted of utilizing a half inch cutter to remove 0.5-mm of material from each side in a single pass. Figures 8 and 9 show the surface of the milled sample and a visualization of how the samples were machined, respectively.1913Figure 8: Image of finished tangential milled sampleFigure 9: Illustration of milling process.The previous methods were performed at UTEP with the milling able to be done at any machine shop. The third machining technique was commissioned to a local shop and consisted of using a grinding wheel in a back-and-forth motion to remove 0.5-mm of material from each side. Figure 11 illustrates how this third process was completed according to the vendor’s process specifications.

Figure 11 illustrates how this third process was completed according to the vendor’s process specifications. Figure 10 demonstrates how the surface of the ground samples looked; scratches can be seen overlapping along the longitudinal axis of the specimen.Figure 10: Image of finished surface ground sampleFigure 11: Illustration of grinding process.The final machining method was also performed by a vendor, Laboratory Testing Incorporated (LTI). The samples were polished and although the company did not provide much information on how the samples were to be polished, it is important to note that the company is an industry certified and accredited site for machining and testing a wide variety of materials. LTI also guaranteed the polished samples have an average line roughness (Ra) of less than or equal to 0.8 microns. Figure 12 below shows the finished surface of these samples observed under a microscope.

Figure 12 below shows the finished surface of these samples observed under a microscope. Scratches can also be seen however these are much more uniform throughout the sample.1914Figure 12: Image of polished sampleSurface quality inspectionLine roughness measurements were taken for all the samples using a mechanical profilometer (Mitutoyo Surftest SJ-210) and optical microscope (Keyence VR-5000) to determine the surface quality that each machining technique produced. The technique that produced the highest roughness values would be considered the worst surface quality and vice versa.Mechanical TestingAn MTS Landmark machine equipped with a 100 kN load cell was utilized to conduct 4- point bending fatigue testing. The testing procedure involved employing two fixtures to establish four-pin connections on the sample, with two pins positioned on the top and two on the bottom.

The testing procedure involved employing two fixtures to establish four-pin connections on the sample, with two pins positioned on the top and two on the bottom. The alignment of the pins ensured an inner span of 10 mm and an outer span of 30 mm on the respective surfaces (loading side) of the sample. The fatigue test was designed to subject the samples to cyclical loading at a stress ratio of 0.1 and a frequency of 10 Hz until failure occurred. To prevent excessive testing duration, a test would be terminated automatically if it reached 7x106 cycles, thereby qualifying as runoff.Fracture Surface AnalysisTo gain a deeper understanding of the failure mechanism exhibited by the samples, the fracture surface analysis was conducted employing a Keyence VHX optical microscope. To facilitate this analysis, the samples underwent a series of isopropanol and acetone baths in an ultrasonic cleaner to effectively eliminate any oils or impurities that might have been present during testing.

Following the cleaning process, the fracture surface of the samples was meticulously examined using the VHX microscope. Multiple images were captured to facilitate the identification of the fracture origin. Additionally, the samples were classified based on the specific location of the fracture, namely surface, chamfer, or internal defect.1915Results and DiscussionAs demonstrated in Table 2 below the average line roughness for all the samples according to their respective surface finishing were recorded. It can be noted that there is a significant improvement in roughness when employing any post processing on the parts. This is illustrated in the vast difference in roughness values between an as built sample and any of the machined samples.

This is illustrated in the vast difference in roughness values between an as built sample and any of the machined samples. Furthermore, it is important to note that when the maximum profile valley depth (Rv) values are compared, the ranking of surface quality from best to worst is: LTI polished, surface ground, tangential milling, Deci Duo, and as built.Table 2: Average Line Roughness for each Machining TechniqueMachining MethodAverage Ra (µm)Average Rv (µm)As Built22.5148.17LTI Polished0.164.544Surface Ground0.4401.737Tangential Milling1.1013.167Deci Duo1.0414.112The raw fatigue data was used to generate figure 13 below comparing the performance of the samples based on the different surface finishes. The data for the machined samples are clustered together with the Deci Duo and milled samples having a slightly better performance over the surface ground and polished samples.

The data for the machined samples are clustered together with the Deci Duo and milled samples having a slightly better performance over the surface ground and polished samples. Additionally, the performance of all machined samples is vastly superior to that of as built samples, as none of the latter ever achieved runoff, even at significantly lower stress levels.1.20Maximum Stress vs Fatigue Lifeh t g n e r t S d e Y t n e c r e Pli1.000.800.600.400.201.E+031.E+041.E+05 Cycles to Failure1.E+061.E+07MilledDeci DuoSurface GroundPolishedAs-BuiltFigure 13: Percent Yield Strength vs. Cycles to Failure Graph1916The graph suggests that the two worst machining methods may improve the fatigue life of the samples, however this cannot be the case. To further understand why the data suggests this, the failure mechanisms of each sample were observed.

To further understand why the data suggests this, the failure mechanisms of each sample were observed. Figures 13, 14, and 15 demonstrate that there were three crack initiation sites: initiation at the surface, at the chamfer, and at an internal defect. Table 3 further suggests that there exists a trend in the fatigue life relative to where the crack initiated. That is that when the crack initiates from the surface, the fatigue life is shorter and when the initiation point is at an internal defect, the fatigue life is longer.

This makes sense; the surface receives stress concentrations much sooner than the chamfer or any internal defect therefore leading to a lower fatigue life.Figure 14: Sample F17 – Initiation site on the lower left side near the chamfer1917Figure 15: Sample F49 - Initiation site on the lower left side chamferFigure 16: Sample F77 - Initiation site at defect on the lower right side1918Table 3: Fracture Initiation Site and Average Fatigue LifeMachining TechniqueInitiation Site SurfaceCount Average Fatigue Life137,852Tangential MillingChamfer548,421Defect41,735,574Surface917,046Surface GroundChamfer1153,446Defect0Surface555,501Deci DuoChamfer0Defect41,185,435Surface414,259LTI PolishedChamfer526,104Defect0However, this correlation between the fracture initiation site and the fatigue life of the corresponding sample, also does not explain the dilemma observed in figure 13.

In fact, the same trend of the specimens with better surface finishes having lower fatigue life is further seen when the data is isolated to only include the samples that failed from the surface.To understand the significance of this relationship two statistical tests were performed on Minitab, a statistical analysis software. The first test was an analysis of variance (ANOVA) amongst the different machining techniques. This was done to see if there was a statistically significant difference between the four processes. Figure 17 demonstrates that the surface ground and polished samples were unique from the other samples and that the milled and Deci Duo samples were like one another due to an overlap in their mean roughness values.Furthermore, the fatigue data and Rv measurements were set to fit a general linear model. The null hypothesis of this test is that there exists no significant correlation between the variables.

The null hypothesis of this test is that there exists no significant correlation between the variables. This hypothesis can be rejected when the P-value of the test is less than 0.05, which indicates a confidence interval of 95 percent. The variables explored for this test were the fatigue life (cycles) versus the machining techniques and the stress levels. This was done to understand if one or both the machining and/or stress levels had a significant impact on the fatigue life of the samples. Figure 18 demonstrates the results of this test: only the testing stress had a significant impact on the fatigue life of the specimens.1919Figure 17: ANOVA results comparing the line roughness values of the different machining techniques.Figure 18: General Linear Model: Cycles versus Machining, Stress results1920ConclusionsIn conclusion, it was observed that machining laser powder bed fusion parts significantly improves fatigue life when comparing the performance of as built samples.

In this study, it was also found that machining technique had no significant correlation to the fatigue life of the specimens, but the rather the maximum testing stress would be the deciding factor when predicting fatigue performance. This could be the case for specimens below a certain line roughness threshold such as in this study where all machined samples had an Rv and Ra of less than 5 and 2 microns, respectively. Another explanation could also be attributed to stresses introduced during the post processing. Further work needs to be completed to understand which is the case. Additionally, it is worth noting that achieving satisfactory results may not require surface quality improvement beyond the capabilities of simple milling techniques.AcknowledgementsThe research described here was performed at The University of Texas at El Paso (UTEP) within the W.M. Keck Center for 3D Innovation (Keck Center).

Keck Center for 3D Innovation (Keck Center). The authors would also like to thank Enrique Garibay and Hunter Taylor for their assistance in the machine setup and the manufacturing of the samples. Primary support for this research was provided by grant number 1110234-424023 from NASA ULI as well as strategic investments via discretionary UTEP Keck Center funds and the Mr. and Mrs. MacIntosh Murchison Chair I in Engineering Endowment at UTEP. The views and opinions expressed in this article are those of the authors and do not necessarily reflect the official opinion or policy of NASA ULI or of any agency of the US government.1921References[1] ASTM F2792-12a, "Standard Terminology for Additive Manufacturing Technologies," ASTM International, 2012[2] N. Guo and M. C. Leu, “Additive manufacturing: technology, applications and research needs,” Front. Mech. Eng., vol. 8, no. 3, pp. 215–243, Sep. 2013, doi: 10.1007/s11465-013-0248- 8.[3] L. Thijs, F. Verhaeghe, T. Craeghs, J. V.

Ednie et al., “The effects of surface finish on the fatigue performance of electron beam melted Ti–6Al–4V,” Materials Science and Engineering: A, vol. 857, p. 144050, Nov. 2022, doi: 10.1016/j.msea.2022.144050.[9] ASTM B215-15, "Standard Practices for Sampling Metal Powders," ASTM International, 2015.[10] ASTM B822-20, "Standard Test Method for Particle Size Distribution of Metal Powders Using Semiautomatic Analytical Electrozone Separation," ASTM International, 20201922.

Size of defects such as gas entrapped pores are captured using fractography analysis and calculated employing Murakami’s approach with the l area method. Considering the value of the area as initial crack length, fatigue life is then calculated using NASGRO software, and compared to the experimental data obtained from strain- controlled fatigue testing. A comparison between predicted fatigue lives using NASGRO software, combined with the Murakami approach, and experimentally obtained ones were presented to determine the applicability of the utilized model for predicting the fatigue performance of additive manufactured materials.lKeywords: Additive manufacturing, defect size, NASGRO, fatigue life prediction.IntroductionAdditive manufacturing (AM), a layer-wise manufacturing method, enables fabricating parts with complex geometries layer by layer [1–3].

The development in AM technologies has revamped a broad spectrum of industries by introducing the potential of manufacturing parts with complex structure, functionally graded materials, and even assemblies, which are often challenging for conventional fabricating processes [4]. Among different AM methods, laser beam powder bed fusion (LB-PBF) is one of the most commonly used. This method uses a laser beam to fabricate parts from a pre-deposited layer of powder [5]. Similar to other AM methods, LB-PBF offers numerous benefits like cost reduction, smaller equipment footprint, reduced labor cost, reduced manufacturing time, no mold and pattern, and most significantly, it could offer enhanced mechanical properties [6]. However, along with numerous advantages, the unique thermal processing history experienced during AM induces many challenges like tensile residual stresses, and the formation of defects like pores, and lack of fusion [1, 6].

Poor surface finish as a result from the repetitious scanning during printing specimens also facilitates crack nucleation. These process-inherent defects are very detrimental when the part is subjected to cyclic loading, as it may lead to premature fatigue failure [2, 7]. In order to address these issues about the structural integrity494of AM parts under cyclic loading conditions, it is crucial to develop a model which can successfully predict fatigue life behavior of additively manufactured materials.17-4 precipitation hardening (PH) stainless steel (SS), being a suitable candidate material for AM, can be widely used to manufacture complex parts with minimum wastage. 17-4 PH SS possesses excellent mechanical properties like high strength, fracture toughness, and corrosion resistance at both ambient and service temperature (below 300°C) after appropriate heat treatment.

Due to the above-mentioned properties, 17-4 PH SS is extensively used in various applications including aviation, power plants, petrochemical, and marine applications [8]. However, when used in the fabrication of additively manufactured parts, inherent residual stresses have proven detrimental to fatigue and mechanical properties [9]; hence, post-process heat treatments have been employed to improve the overall mechanical performance of AM 17-4 PH SS. Therefore, it is critical to address the specific type of heat treatment condition, its effects on the fatigue behavior of 17-4 PH SS, and incorporate them when attempting to use fatigue life prediction model [10].Until now, numerous researchers followed different approaches to model fatigue life behavior of additive manufactured materials [12, 13]. Torries et al., to calculate the fatigue life of AM Ti-6Al-4V, employed microstructure sensitive fatigue (MSF) model [14–16]. Beretta et al. [17], and Romano et al.

Torries et al., to calculate the fatigue life of AM Ti-6Al-4V, employed microstructure sensitive fatigue (MSF) model [14–16]. Beretta et al. [17], and Romano et al. [18, 19] adopted ‘defect tolerant design’ concept based on ‘Kitagawa- Takahashi’ diagram to model fatigue behavior of additive manufactured materials. NASGRO is a well-known fatigue behavior analysis tool based on fracture mechanics approach [20]. Murakami’s (cid:958)(cid:1853)(cid:1870)(cid:1857)(cid:1853) method [21, 22] is one of the promising approaches to measure the crack size based on the projection of the defect size on the loading plane in the loading direction, which can be incorporated in NASGRO crack growth equation to calculate fatigue life.

However, to the best of authors knowledge, based on Murakami’s approach, using the NASGRO crack propagation equation to assess the fatigue life performance of the LB-PBF 17-4 PH SS after CA-H900 heat treatment has not been studied hitherto.Here, this study aimed to develop a defect sensitive fatigue model to predict the fatigue life of LB-PBF 17-4 PH SS using fracture mechanics approach based on the known defect population of the material.

As an effort to predict the fatigue life, NASGRO software was used to calculate fatigue life using the inherent defect size of the materials based on Murakami’s approach, and a comparison between NASGRO predicted fatigue lives and experimentally obtained values were presented to determine the applicability of this model for predicting the fatigue performance of AM materials.Material and Specimen FabricationUsing argon-atomized 17-4 PH SS powder, the samples were prepared in EOS M290 additive manufacturing machine using standard process parameter. Square bar and net-shape round specimens were fabricated and fatigue tests were conducted after employing CA-H900 heat treatment [23] on the LB-PBF 17-4 PH SS specimens. Square bars (11 mm X 11mm X 77 mm) were machined to the uniform round specimen and both the as-built and machined specimens have the similar final dimensions of 15 mm uniform gauge length with 5 mm gauge diameter.

For CA- H900 heat treatment, samples were heated at 1050°C for 30 minutes followed by air-cooling. Then, the samples were heated at 482°C for 1 hour. Detailed results are reported in [23]. The experimental495fatigue life data used in this investigation was obtained by conducting strain-controlled uniaxial fully reversed (R(cid:4242) = -1) fatigue test in MTS landmark servohydraulic testing machine based on ASTM E606 standard [24] at ambient temperature. Fatigue crack growth data for LB-PBF 17-4 PH SS were adapted from [25]. To analyze defect sizes for LB-PBF 17-4 PH SS, a Keyence VHX- 6000 microscope was used.Defect Sensitive Fatigue ModelIn general, there are three stages of fatigue life in a defect sensitive material: crack nucleation, crack propagation, and final fracture. When the material undergoes cyclic loading below the yield stress, crack nucleation occurs. The nucleation of the crack is typically attributed to the slip mechanisms.

When the material undergoes cyclic loading below the yield stress, crack nucleation occurs. The nucleation of the crack is typically attributed to the slip mechanisms. With the continuation of the loading cycles, the crack advances with the plastic deformation at the crack tip. When the crack size reaches the final fracture zone, i.e., critical crack size, final fracture occurs [26, 27]. However, with the current additive manufacturing technology, it is impossible to fabricate a defect-free part due to the formation of pores [1, 28]. These defects can act as a source for initial crack nucleation. These defects can be considered as pre-existing cracks playing the role of internal stress raiser. Therefore, the fatigue life of AM materials, in general, is mainly spent in the second stage, i.e., the crack propagation phase [15, 19, 29–31].

Therefore, the fatigue life of AM materials, in general, is mainly spent in the second stage, i.e., the crack propagation phase [15, 19, 29–31]. Therefore, to appropriately evaluate the fatigue life of AM parts, it is critical to consider AM inherent defects as pre-existing cracks during the crack propagation phase. Currently, there are three approaches for calculating fatigue life considering these structural features: fracture mechanics, the multistage fatigue model (MSF), and numerical modeling [7].Stress intensity factor, K, indicates the state of unstable or rapid crack propagation without any increment of applied stress.

Murakami’s (cid:958)(cid:1853)(cid:1870)(cid:1857)(cid:1853) approach is a well-established method for the estimation of effective size of irregular defects and interacting adjacent defects for fatigue design [32], and has proved reliable when used in fatigue life calculations [21, 22]. In this approach, a threshold stress at the tip of cracks developing from inherent defects, e.g. pores or LoFs, is determined to calculate fatigue life of the specimen by employing equation (1). The initial defect size was estimated by measuring the square root of the defect area projected in the direction of applied loading and then the value was input into the equation.

Crack is nucleated in the crack nucleation zone, and this zone is mainly controlled by factors like microstructure, mean stress, etc. Crack propagation zone depends on the relationship between the crack size corresponding to the number of cycles to the stress intensity factor [33].(cid:1856)(cid:1853) (cid:1856)(cid:1840)= (cid:1829)((cid:1986)(cid:1837))(cid:3041) (4)Here, a is the crack length, N is the number of cycles, C is a intercept constant, and n is theslope of crack growth rate and stress intensity plotting on a log-log scale.After obtaining the value of stress intensity factor, K, it is feasible to get the fatigue crack growth life by integrating the sigmoidal curve of fatigue behavior between the limits of initial defect size and final defect size. However, employing equation (4) is not enough to accurately calculate the correct fatigue behavior of materials as it is not capable to accounts for stress ratio.

NASGRO software also considered the effect of crack closure while calculating the fatigue life. In this research, only the life of the fatigue crack propagation stage was calculated, and the values of p and q were assumed as 0. Therefore, the equation used in this study is:(cid:1856)(cid:1853) (cid:1856)(cid:1840)= (cid:1829) (cid:3428)(cid:3436)1 (cid:3398) (cid:1858) 1 (cid:3398) (cid:1844)(cid:3440) (cid:1986)(cid:1837)(cid:3432)(cid:3041)(6)497Results and DiscussionFigure 1: Crack initiation from defects (pore) for machined LB-PBF 17-4 PH SS after CA-H900 heat treatmentFigure 2: Crack initiation from the micro-notches of the boundary surface for as-built LB-PBF 17-4 PH SS after CA-H900 heat treatmentFigure 1 and 2 demonstrate the crack initiation site of machined and as-built LB-PBF 17- 4 PH SS, respectively. The crack nucleation sites are marked with red arrows and, for machined specimens, the crack sources are marked with red circle.

The crack nucleation sites are marked with red arrows and, for machined specimens, the crack sources are marked with red circle. For machined specimens, the source of crack initiation is uniform round pores. According to the Murakami’s approach, size of the pores was analyzed to calculate the defect size [21]. On the other hand, as-built specimens have multiple micro-notches on the surface due to poor surface finish attributed to the additive manufacturing process, which might act as a source of crack initiation. These micro-notches make it challenging to measure defect size accurately.

These micro-notches make it challenging to measure defect size accurately. Hence, for AM parts, considering the post-contour process, a scanning strategy after melting the powder at the edge of the surface for better surface finish, a conservative approach may be followed to determine the crack size of as-built parts [34]:498(cid:3493)(cid:1853)(cid:1870)(cid:1857)(cid:1853)(cid:3028)(cid:3046)(cid:2879)(cid:3029)(cid:3048)(cid:3036)(cid:3039)(cid:3047) =(cid:3557) (cid:1856)(cid:958)10 (7)Here, d is the width of the area covering the post-contour process on the circular fracture surface of the specimen, which is typically within the range of 60-80 μm for LB-PBF process employing EOS machine.Figure 3 demonstrates the fatigue life correlations between the NASGRO predictions and experimental fatigue life of LB-PBF 17-4 PH SS in both as-built and machined conditions after CA-H900 heat treatment.

In Figure 3, blue line demonstrates the perfect fit line and orange dashed lines demonstrate the scatter band of factor of 2. Scatter bands of a factor of 2 are also included for data comparison. Fatigue life was calculated according to the equation (6) and (7), utilizing fracture mechanics approach. Although CA-H900 procedure of LB-PBF 17-4 PH SS demonstrates superior mechanical properties (such as high strength), its fatigue performance is poor. Due to the high strength obtained after CA-H900 procedure, the material is highly sensitive to defects such as LoF, which decreases the fatigue resistance of the material specifically in high cycle fatigue regime [35]. From Figure 3, it is observed that 100% of the predicted fatigue data are within the scatter band of factor of 2 for both as-built and machined LB-PBF 17-4 PH SS specimens. For machined specimens, mostly a non-conservative prediction is observed for both low cycle fatigue and high cycle fatigue regime.

For machined specimens, mostly a non-conservative prediction is observed for both low cycle fatigue and high cycle fatigue regime. For as-built specimens, mostly conservative predictions were observed._ 1E+07 z C'\IMachined specimens~ -; 1E+06 ::::, C, .:; ~ "C 1E+05a, -(J a, ... a.1E+04 0 0::: (!) UJ ~ 1E+03"C♦ As-built specimens1E+031E+04 1E+05 1E+06 Experimental fatigue life, 2N,1E+07Figure 2: NASGRO calculated fatigue life correlation to the experimental fatigue life of both as-built and machined LB-PBF 17-4 PH SS after CA-H900 heat treatment499ConclusionsIn this study, the fatigue life prediction of LB-PBF 17-4 PH SS was investigated based on fracture mechanics approach employing Murakami’s (cid:958)(cid:1853)(cid:1870)(cid:1857)(cid:1853) method.

This investigation can be summarized as follows:•• NASGRO crack growth equation incorporating Murakami’s (cid:958)(cid:1853)(cid:1870)(cid:1857)(cid:1853) approach is capable to predict the fatigue life behavior of both as-built and machined LB-PBF 17-4PH SS for CA- H900 heat treatment conditions.Fatigue life predictions obtained using NASGRO, for both as-built and machined LB-PBF 17-4 PH SS specimens, fall within scatter bands of factor of 2, which shows that the estimated fatigue life is in good agreement with the experimental fatigue life.A non-conservative fatigue life prediction is obtained for machined LB-PBF 17-4 PH SS for CA-H900 heat treatment conditions which might be attributed to the better surface finish of machined samples.A conservative fatigue life prediction is obtained for as-built LB-PBF 17-4 PH SS for CA- H900 heat treatment conditions which is accounted to the conservative approach employed due to poor surface finish.AcknowledgmentThe authors greatly acknowledge for the support provided by the National ScienceFoundation under grant #1657195.References[1] N.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFatigue Life Prediction of Additively Manufactured Metallic Materials Using a Fracture Mechanics ApproachBrian Torries1,2, Rakish Shrestha1,2, Aidin Imandoust1,2, Nima Shamsaei1,2 *1Department of Mechanical Engineering, Auburn University, Auburn, AL 368492National Center for Additive Manufacturing Excellence (NCAME), Auburn University,Auburn, AL 36849Corresponding author: shamsaei@auburn.eduAbstractThe present study aims to model the fatigue strength of additively manufactured metallic materials employing a fracture mechanics approach. Specimens with different build orientations were subjected to strain controlled fatigue testing. Upon failure, the defect(s) responsible for crack initiation were identified by fractographic analysis. From these defects an equivalent internal method based on Murakami model.

Upon failure, the defect(s) responsible for crack initiation were identified by fractographic analysis. From these defects an equivalent internal method based on Murakami model. Using this parameter, defect size is calculated using the ) was determined, and the relationship between the elastic-plastic energy release rate ( and fatigue life was investigated. The results showed that this method improves the predictability 𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒 of the fatigue strength of additively manufactured materials when the defects size and location is known. The relationship appeared to better fit the fatigue data of the experimental materials as compared to the√𝑎𝑎𝑎𝑎𝑎𝑎𝑎𝑎𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒 − 𝑁𝑁𝑒𝑒relationship and contributed to a reduction in data scatter.𝜀𝜀𝑎𝑎 − 𝑁𝑁𝑒𝑒Keywords: Additive manufacturing, Defects, Fracture mechanics, Fatigue modelingIntroductionAdditive manufacturing (AM) is an attractive fabrication method with many benefits.

AM allows for the fabrication of parts with complex geometries unobtainable through traditional machining processes [1-3], functionally graded materials [4-9], and internal porosity [4, 6, 9]. Additionally, AM can be used to fabricate parts customized for specific applications. However, despite these advantages, AM cannot attain full industry adoption until the effects of structural inconsistencies on the tensile and fatigue properties of additively manufactured components can be accounted for.From a fatigue perspective, some of the most impactful structural inconsistencies found in additively manufactured components are the defects, such as pores and lack of fusion (LOF) voids, inherent to the AM process. These internal defects can act as stress risers and cause premature fatigue failure of the specimens [10, 11]. Pores and LOF have also been shown to contribute to the scatter found in the fatigue data of additively manufactured materials [12, 13].

Pores and LOF have also been shown to contribute to the scatter found in the fatigue data of additively manufactured materials [12, 13]. While post processing techniques such as hot isostatic pressing (HIP) can help shrink these pores and improve fatigue performance [14-16], its application might not be desirable due to the changes in microstructure stemming from the applied heat and pressure. Additionally, HIP has been shown to1181be less effective at improving the fatigue properties of specimens with surfaces in the as-built condition, as the detrimental effects of surface roughness may override those of internal defects [16-19]. At the current state of AM technology, it is nearly impossible to fabricate defect free additively manufactured specimens.

At the current state of AM technology, it is nearly impossible to fabricate defect free additively manufactured specimens. Therefore, a defect sensitive fatigue analysis is required to further improve the predictability of the fatigue response of AM parts.Multiple models have been used to account for the fatigue behavior of additively manufactured materials, each accounting for different sources of uncertainty [10, 20-24]. One model that has exhibited promising results in accounting for the effects of process inherent defects is the model by Murakami et al. [25, 28]. This model uses a stress intensity factor calculated via an √𝑎𝑎𝑎𝑎𝑎𝑎𝑎𝑎 equivalent internal defect size derived from the defect area projected on the loading plane, while also accounting for the location of the defect.In this study, the fatigue behavior of 316L stainless steel (SS) and Ti-6Al-4V are investigated.

Fractography analysis was used to identify the defects responsible for crack initiation and failure of the specimens. From these defects, the equivalent internal flaw sizes were determined using the method developed by Murakami et al [25]. Using the equivalent flaw size, the elastic- was calculated, and the relationship between defect size andplastic energy release rate √𝑎𝑎𝑎𝑎𝑎𝑎𝑎𝑎 fatigue life was then investigated. 𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒Experimental ProcedureThe strain-life data utilized in this study was obtained from strain-controlled fully reversed (Rε = -1) fatigue tests conducted on both additively manufactured 316L SS and Ti-6Al-4V, reported in detail in [29]and [30], respectively. The specimens were fabricated in three different build orientations (vertical, diagonal (45°), and horizontal) via laser based powder bed fusion (L-PBF) process in an argon environment.

The specimens were fabricated in three different build orientations (vertical, diagonal (45°), and horizontal) via laser based powder bed fusion (L-PBF) process in an argon environment. The 316L SS specimens were fabricated via a Renishaw AM250 machine [29], while the Ti-6Al-4V specimens were fabricated utilizing an EOS M290 [28].After fatigue testing, the fracture surfaces were carefully examined using a scanning electron microscope (SEM) in order to determine the factors responsible for crack initiation and the failure mechanism in each of the material. Special care was taken to determine if multiple crack initiation sites were present on each fracture surface. If multiple initiation sites were found, the largest one was used in thecalculations.𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒Fracture MechanicsIn defect free materials, the fatigue life can be divided into three stages: crack initiation, crack growth, and final fracture.

First, cyclic loading and plastic deformation causes the movement of dislocations within the material, called “slip.” A crack then forms due to formation of persistent slip bands or intrusions and extrusions on the specimen surface [31], or due to slip between grain boundaries [30]. This crack grows due to plastic deformation at the crack tip stemming from cyclic loading and unloading, and, once the crack reaches critical size, final fracture occurs [31]. However, additively manufactured materials can behave differently than typically processed materials. During AM, pores and LOF voids can form due to sub-optimal process parameters [33, 34], high speed flow [35] or vapor recoil [36] in the melt-pool, or from porosity contained in the feedstock used during fabrication [12, 33]. These process inherent defects can be considered as pre-existing cracks, thereby shortening or bypassing the crack initiation phase [12, 15, 20, 22, 37, 38].

These process inherent defects can be considered as pre-existing cracks, thereby shortening or bypassing the crack initiation phase [12, 15, 20, 22, 37, 38]. Fractography analysis of the specimens used in this study revealed pores from entrapped1182gases, as well as LOF and defects resulting from partially melted powder particles. These defects were present at random locations and varied in both shape and size, as can be seen in Fig. 1(a-b) and 1(c-d) in 316L SS and Ti-6Al-4V specimens, respectively.Fig. 1.

calculated using [26].is the stress range, and𝑛𝑛′(1)𝛥𝛥𝛥𝛥𝑒𝑒𝑒𝑒𝑒𝑒𝛥𝛥𝜀𝜀𝑃𝑃𝑃𝑃,𝑒𝑒𝑒𝑒𝑒𝑒(2)𝛥𝛥𝛥𝛥𝑒𝑒𝑒𝑒𝑒𝑒 = 𝛥𝛥𝑚𝑚𝑎𝑎𝑚𝑚 − 𝛥𝛥𝑐𝑐𝑃𝑃(3)𝛥𝛥𝜀𝜀𝑝𝑝𝑃𝑃,𝑒𝑒𝑒𝑒𝑒𝑒 = 𝛥𝛥𝜀𝜀𝑒𝑒𝑒𝑒𝑒𝑒 − 𝛥𝛥𝜀𝜀𝑒𝑒,𝑒𝑒𝑒𝑒𝑒𝑒where induced crack closure theory presented in [26, 27]. Similarly, calculated by subtracting crack tip opening strain from the maximum strain, andis the maximums stress,is the crack closing stress that is related to the plasticity is the effective strain range is the𝛥𝛥𝑚𝑚𝑎𝑎𝑚𝑚𝛥𝛥𝑐𝑐𝑃𝑃𝛥𝛥𝜀𝜀𝑒𝑒𝑒𝑒𝑒𝑒𝛥𝛥𝜀𝜀𝑒𝑒,𝑒𝑒𝑒𝑒𝑒𝑒1183elastic portion of the effective strain range defined as the ratio of cyclic modulus and effective stress range [26]:𝛥𝛥𝜀𝜀𝑒𝑒𝑒𝑒𝑒𝑒 = 𝜀𝜀𝑚𝑚𝑎𝑎𝑚𝑚 − 𝜀𝜀𝑜𝑜𝑝𝑝𝐸𝐸′ ∆𝛥𝛥𝑒𝑒𝑒𝑒𝑒𝑒∆𝜀𝜀𝑒𝑒,𝑒𝑒𝑒𝑒𝑒𝑒 =in Eq. 1 accounts for both the plastic and elastic energy released It should be noted that during the life of a component. This is important, as at high strain amplitudes, the fatigue deformation is governed by both elastic and plastic strains.

This is important, as at high strain amplitudes, the fatigue deformation is governed by both elastic and plastic strains. At lower strain amplitudes, the fatigue deformation was almost entirely governed by elastic strain ( can be related to stress intensity factor,𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒0). In such cases,, as [31]:𝛥𝛥𝜀𝜀𝑃𝑃𝑃𝑃,𝑒𝑒𝑒𝑒𝑒𝑒 ≈𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒𝐾𝐾𝑚𝑚𝑎𝑎𝑚𝑚Note that the as shown in Fig. 1.�𝑎𝑎𝑎𝑎𝑎𝑎𝑎𝑎𝑒𝑒𝑒𝑒𝑒𝑒′used in these calculations can be greater than the actual size of the defect,𝐾𝐾𝑚𝑚𝑎𝑎𝑚𝑚 = �𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒 𝐸𝐸Results and DiscussionThe strain life correlations for L-PBF 316L SS and Ti-6Al-4V are shown in Fig. 2(a) and 2(b), respectively, along with factor of two and factor of three scatter bands. Relatively larger amounts of scatter in the fatigue life data of both 316L SS and Ti-6Al-4V were primarily observed in the high cycle fatigue (HCF) regime, where the influence of porosity and other defects on fatigue life is more pronounced [12, 34].

For 316L SS, the overall larger scatter was attributed to the inconsistent variations in size and distribution of LOF defects serving as crack initiation factors in the tested specimens [29]. On the other hand, for the Ti-6Al-4V specimens, the scatter in fatigue life data was attributed to more uniformly shaped pores serving as crack initiation sites rather than LOF as seen in the 316L SS [30]. Additionally, as was noted in the initial study of the 316L SS specimens, due to the influence of LOF defect orientation with respect to the loading direction there were some build orientation effects observed in the 316L SS specimens [29]. Results showed that horizontally fabricated specimens exhibited higher fatigue life as compared to the vertical specimens, which in turn performed better than the diagonal specimens [29]. Furthermore, from Fig.

Furthermore, from Fig. 2, the inability of strain amplitude to correlate the fatigue life data was also evident as only 73.7% and 37.5% of the fatigue life data were within the factor of two scatter band for 316L SS and Ti-6Al-4V, respectively. Therefore, because the presence of defects such as LOF and gas entrapped pores were considered as the factor responsible for this scatter in the data, a damage parameter that can incorporate the effects of defects might be able to accurately correlate the fatigue life data for both AM materials.1184(4)(5)(2)Scatter Band of 2, 73.7% Scatter Band of 3, 94.7%(a)Scatter Band of 2, 73.7% Scatter Band of 3, 94.7%(b)Fig. 2. Strain versus reversals to failure for fully-reversed, strain-controlled fatigue tests of machined L-PBF (a) 316L SS [29] and (b) Ti-6Al-4V [30] specimens.was calculated using Eq. 1. The versus fatigue life correlation, plotted in Fig.

1. The versus fatigue life correlation, plotted in Fig. 3, was shown to satisfactorily model the fatigue behavior of both the L-PBF Ti-6Al-4V and 316L SS. The scatter observed in the strain-life relationship appears to have been reduced, as 84% and 81.3% of the fatigue data were within the factor of 2 scatter bands as compared to 73.8% and 37.5% for 316L SS and Ti-6Al-4V, respectively. It is interesting to note that the amount of data points falling within the factor of 3 scatter band did not change for the 316L SS, but did increase for the Ti-6Al- relationship was able to fit the curve well due to 4V, increasing from 75.0% to 93.7%. The a number of factors. has been used to assess fatigue life in low cycle fatigue (LCF) scenarios in which cracks begin propagating during the first loading cycle [39].

The a number of factors. has been used to assess fatigue life in low cycle fatigue (LCF) scenarios in which cracks begin propagating during the first loading cycle [39]. As mentioned previously,After the fractographic investigation andcalculations were completed,�𝑎𝑎𝑎𝑎𝑎𝑎𝑎𝑎𝑒𝑒𝑒𝑒𝑒𝑒𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒1185the defects such as voids and pores present in additively manufactured materials can be treated as preexisting cracks [12, 15, 20, 22, 37, 38]. Therefore, crack propagation can be considered to begin in the first loading cycle in additively manufactured materials as well. Furthermore, calculation of also considers the local plastic deformation in crack tip by incorporating effective stress, parameter and shape which is based on the crack closer concept. Through the use of the 𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒 factor Y, can also account for defect size and location.

Through the use of the 𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒 factor Y, can also account for defect size and location. Defects have been shown to be a source of scatter in strain-controlled fatigue data for additively manufactured components; therefore, accounting for the effects of size and location of such internal defects is vital to accurately correlate the fatigue life data.�𝑎𝑎𝑎𝑎𝑎𝑎𝑎𝑎𝑒𝑒𝑒𝑒𝑒𝑒𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒Scatter Band of 2, 84.2% Scatter Band of 3, 94.7%(a)Scatter Band of 2, 81.3% Scatter Band of 3, 93.7%(b) Fig. 3. ΔJeff-life relationships for L-PBF (a) 316L SS [29] and (b) Ti-6Al-4V [30] specimens.Conclusions1186was calculated for L-PBF 316L stainless steel and Ti-6Al-4V specimens fabricated in multiple build orientations, and its relationship to fatigue life was investigated based on the type, size, shape, and location of the defects responsible for the failure in the specimens. From these results, the following conclusions have been drawn:In this study, the𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒1.

From these results, the following conclusions have been drawn:In this study, the𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒1. Alife relationship satisfactorily fit the fatigue behavior of both the 316L SS and Ti-6Al-4V L-PBF specimens.𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒2. Thelife relationship for both materials exhibited reduced scatter as compared to thecorresponding strain-life data.𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒3. The better correlation of fatigue life data obtained usingas a damage parameter can be attributed to its ability to incorporate shape, size, and location of the defects along with localized plastic deformation at the tip of the these defects, which acts as a small crack.𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒 may be utilized to predict fatigue4. Therefore, with the crack growth properties known, life in AM materials using crack growth behavior 𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒can account for the internal defects inherent to the additive process and reduce the scatter in the fatigue response of additively manufactured materials.

Defect shape, however, may be a factor in the effectiveness of this approach. Further study is needed to investigate the effects of the defect geometry serving as the crack initiation site, as well as the applicability to other material systems and specimen conditions.From the results of this study, it has been shown that𝛥𝛥𝐽𝐽𝑒𝑒𝑒𝑒𝑒𝑒Acknowledgement This material is based upon work partially supported by the National Science Foundation (NSF) under Grant No. 1657195 and the National Aeronautics and Space Administration (NASA) Award No. NNX17AJ93A and 80MSFC17M0023.1187References [1] D. S. Thomas and S. W. Gilbert, “Costs and Cost Effectiveness of Additive Manufacturing: A Literature Review and Discussion.” National Institute of Standards and Technology, 2014.[2] D. S. Thomas, “Costs, Benefits, and Adoption of Additive Manufacturing: A Supply ChainPerspective,” Int. J. Adv. Manuf., vol. 85, no. 5–8, pp. 1857–1876, 2016.[3] K. S. Chan, M. Koike, R. L. Mason, and T.

Current research studies demonstrate the potential of functionally graded materials (FGMs) in enhancing the mechanical properties of engineered structures. The raw materials employed for the experimentation of this study are a combination of Polylactic acid (PLA) and Thermoplastic Polyurethane (TPU). To predict fatigue life, several researchers have utilized various statistical approaches. In this investigation, an experimental study is conducted utilizing Tension-Tension (T-T) loading conditions and different stress levels (80, 60, 40, and 20% of Ultimate tensile strength), followed by the application of Basquin’s Model for fatigue life prediction. The results obtained indicate that the model may be utilized to predict fatigue response.

The results obtained indicate that the model may be utilized to predict fatigue response. Overall, the soft-hard material combinations with adaptable properties produced through FGAM have potential applications in dental and orthopedic fields.Keywords: Fatigue life; Fatigue Prediction; Basquin’s Model; S-N Curve; FGAMIntroductionMulti-material additive manufacturing (MMAM) has engrossed numerous researchers due to its inherent advantages. MMAM facilitates the fabrication of parts encompassing diverse materials in a single manufacturing process, wherein these materials can exhibit distinct chemical, physical, mechanical, and electrical properties [1]. In the last twenty years, AM has emerged as a pivotal technology within the manufacturing industry. AM technologies have revolutionized traditional manufacturing methods by offering several significant advantages [2].

AM technologies have revolutionized traditional manufacturing methods by offering several significant advantages [2]. Compared to conventional approaches, AM eliminates the requirement for extensive tooling [3], resulting in cost and time savings [4]. Additionally, AM provides unparalleled flexibility in design and allows for easy modification of products during the manufacturing process. These capabilities empower manufacturers to rapidly iterate and customize designs, leading to improved efficiency and innovation in the production of various components and products [5]. MMAM encompasses a comprehensive classification comprising seven distinct technologies. These include material extrusion, vat photopolymerization, powder bed fusion, material jetting, direct energy deposition, sheet lamination, binder jetting, and hybrid additive manufacturing [6].

These include material extrusion, vat photopolymerization, powder bed fusion, material jetting, direct energy deposition, sheet lamination, binder jetting, and hybrid additive manufacturing [6]. Each technology offers unique capabilities and characteristics, contributing to the versatility and potential of MMAM in achieving precise material combinations and complex part geometries [7]. The MEX process in MMAM, a variety of materials are utilized, including thermoplastics, metal-filled thermoplastics, composites, and more. Figure 1 provides a schematic view depicting the process description of the MEX process in MMAM.68Figure 1: Process description of MEX process [8]FGMs represent cutting-edge composite materials that exhibit gradual variation in constituent composition and/or microstructure along specific directions [9]. This unique characteristic distinguishes FGMs from traditionally manufactured laminated composites and offers a significant advantage.

This unique characteristic distinguishes FGMs from traditionally manufactured laminated composites and offers a significant advantage. The smooth and continuous variation in material properties within FGMs mitigates the occurrence of debonding and delamination failures commonly associated with laminates [10]. The customized nature of FGMs allows for the attainment of superior qualities in thermal conductivity, hardness, mechanical strength, and electrical conductivity. By tailoring the material properties to specific requirements, FGMs enable enhanced performance and durability in a wide range of applications [11]. FGMs can be produced using various manufacturing technologies, but AM presents a compelling opportunity to address the limitations of conventional manufacturing processes for composite materials.

FGMs can be produced using various manufacturing technologies, but AM presents a compelling opportunity to address the limitations of conventional manufacturing processes for composite materials. AM enables the simultaneous deposition of multiple materials in pre-defined ratios, resulting in a controlled variation of composition and microstructure across different regions of the fabricated component. This ability to customize material characteristics in specific areas offers a significant advantage in tailoring the properties of FGMs according to desired specifications. By leveraging AM, manufacturers can achieve precise control over material composition, enabling the creation of complex and highly functional components with enhanced performance and unique material characteristics [10]–[12].Thermoplastic materials exhibit inherent anisotropic characteristics [13].

For several decades, the S-N behavior has been fundamental in the analysis of material fatigue life, tracing its roots back to the nineteenth century [14]. Characterizing the S-N curve holds great importance not only for engineering materials but also for accurately predicting the fatigue life of various components under cyclic loads. However, the prediction of fatigue life remains a challenging task due to the complex nature of fatigue mechanisms and the numerous factors that influence the fatigue behavior of materials and structures [15]. Furthermore, thermoplastics exhibits a wide range of damage mechanisms. In the case of composite materials, their fatigue behavior differs significantly from isotropic materials like metals. Micro-cracks in composites can initiate at an early stage of loading, yet the materials can still endure the load until ultimate failure occurs.

Micro-cracks in composites can initiate at an early stage of loading, yet the materials can still endure the load until ultimate failure occurs. However, to establish fatigue S-N curves, ASTM, and ISO standards [16] define experimental parameters and procedures that involve testing69multiple samples. Moreover, to minimize costs in terms of time and resources, it is crucial to explore statistical methods for curve fitting. Various models, such as Basquin's model [17], the Weibull distribution approach [18], the Coffin-Manson relation [19], and advanced statistical analyses like artificial neural networks (ANN) [20] , have been successfully employed to predict and fit the fatigue behavior of diverse materials. Another one-of-a-kind study was conducted by Eskandari et al. [19] proposed a framework for validating a fatigue damage function and developed a theory for predicting the remaining fatigue life at different stress levels with a constant stress ratio.

[19] proposed a framework for validating a fatigue damage function and developed a theory for predicting the remaining fatigue life at different stress levels with a constant stress ratio. They experimentally validated the theory, focusing solely on the zero-stress condition. Later, Pertuz et al. [21] investigated the effects of various parameters, including fiber-reinforcing materials (fiberglass, Kevlar, and carbon fiber), infill pattern (triangular and hexagonal), infill orientation (0°, 45°, and 60°), and concentric ring reinforcements. The researchers utilized Basquin's model for predicting fatigue life due to its simplicity. They established a logarithmic relationship between stress (S) and cycles to failure (Nf) for loading conditions corresponding to 95%, 90%, 85%, and 80% of the ultimate tensile stress. To analyze the obtained data at each load level, the Weibull distribution was selected as the statistical method for data adjustment.

To analyze the obtained data at each load level, the Weibull distribution was selected as the statistical method for data adjustment. Another one-of-a-kind study conducted by Burhan et al. [13] employed several statistical models to predict fatigue life in the study. The analysis included a broad range of stress ratios, with a focus on reviewing and analyzing S-N curve models. The findings revealed that the models developed specifically for characterizing fatigue data demonstrated significantly better data fitting capabilities compared to the models designed to forecast the impact of stress ratios. Additionally, Marino et al. [22] focused on examining the fatigue properties of different polymer materials. The experimental data was validated using the mathematical Basquin's model. Specifically, the research work highlighted the rotating bending fatigue data obtained from 3D printed specimens fabricated using the specified materials.

Specifically, the research work highlighted the rotating bending fatigue data obtained from 3D printed specimens fabricated using the specified materials. The specimens were subjected to various loading conditions for the purpose of analysis. Similarly, Ghods et al. [23] conducted fatigue life study on Grade 5 Ti6Al4V fabricated through electron beam powder bed fusion. The study investigated the influence of various build and component design parameters on fatigue performance. The fatigue responses were used to construct stress-life diagrams, and the data was fitted to Basquin's model. The stress-life distributions were then statistically analyzed using multiple linear regression to evaluate their relationship with the factors under consideration. Eventually, Mu et al. [24] introduced a novel S-N model, to elucidate the relationships between loads and fatigue life in the context of cyclic loading with constant amplitude.

Eventually, Mu et al. [24] introduced a novel S-N model, to elucidate the relationships between loads and fatigue life in the context of cyclic loading with constant amplitude. This model offers the ability to assess the fatigue life behavior of fiber reinforced plastic composites under various loading conditions, including tension-tension, tension-compression, or compression-compression loading, encompassing a range of fatigue life stress ratios. Pertuz et al. [21] conducted fatigue bending tests on continuous fiber-reinforced composites fabricated using the MEX technique under different loading conditions. The composites consisted of Onyx as the matrix material with 19% continuous Kevlar fiber reinforcement. The experimental data obtained from the tests were fitted to Basquin's model, which accurately represented 95.2% of the experimental data. Analysis of the parameter "b" in Basquin's equation indicates that the composite exhibits brittle behavior.

Analysis of the parameter "b" in Basquin's equation indicates that the composite exhibits brittle behavior. Finally, Fard et al.[25] performed fatigue tests on samples, both with and without notches, utilizing a load ratio of 0.1. The experimental results were then compared to predictions. The findings indicated that the fatigue strength reduction factors and fatigue lifetimes were notably influenced by the raster orientation parameter. To date, no published studies have focused on predicting the fatigue behavior of FGMs produced using MMAM. Thus, the objective of this paper is to validate the experimental fatigue data using Basquin's model. A statistical model is employed to forecast the fatigue life of FGM70components manufactured using TPU and PLA. The experimental data is utilized to determine the coefficients of Basquin's model and predict the fatigue behavior. Subsequently, the predicted values by Basquin's model are benchmarked against the remaining experimental data.

Subsequently, the predicted values by Basquin's model are benchmarked against the remaining experimental data. The research work makes a unique contribution through the design, manufacturing, and characterization of additively manufactured components using the MMAM process. Figure 2 illustrates the process for obtaining fatigue life predictions for FGM test samples. The specimen preparation precedes the tensile and fatigue testing, and the generated experimental data is then employed for fatigue life prediction.Figure 2: Flowchart for statistical analysis of fatigue life predictionMaterials and MethodsThe present study employed amorphous polymers such as TPU and PLA. These polymers were obtained commercially in filament form from the Push Plastic company.

These polymers were obtained commercially in filament form from the Push Plastic company. Table 1 presents the parameters that were utilized to 3D print the specimens in the investigation.Table 1: Fixed processing parametersParametersValuesNozzle temperature (oC)210Bed temperature (oC)60Infill density (%)100Infill patternLine (0/90)71Layer width (mm)0.35Layer height (mm)0.2Printing speed (mm/sec)15TPU exhibits excellent flexibility throughout a wide range of hardness levels, making it highly versatile. It possesses exceptional low-temperature performance and impact strength. TPU also demonstrates excellent resistance to abrasion, oil, and grease. The ideal conditions for printing TPU specimens such as bed temperature, nozzle temperature, printing speed are 60°C, 220°C, and 15 mm/s are recommended [11]. Notably, it is crucial to disable retraction settings for optimal results.

Notably, it is crucial to disable retraction settings for optimal results. TPU finds extensive applications across various industries, including automotive[26], aerospace, mobile coverings, medical equipment, and wire and cable manufacturing, owing to its unique combination of plastic and rubber properties [27]. PLA is a thermoplastic polymer renowned for its biodegradability and absorbability. It is frequently employed in rapid prototyping applications. PLA exhibits excellent strength and stiffness at room temperature [28]. It is commonly used in the production of biodegradable medical devices like stents and implanted medication dispensers. Additionally, PLA is utilized in food packaging and disposable flatware due to its favorable properties. With its low melt temperature and user-friendly nature [29], PLA is one of the most widely used filaments for the MEX 3D printing process[30].

With its low melt temperature and user-friendly nature [29], PLA is one of the most widely used filaments for the MEX 3D printing process[30]. Table 2 provides an overview of the material properties of TPU and PLA polymers.Table 2: Properties of TPU and PLAParametersValuesTPUPLAMelting point (oC)190-210200-230Density (kg/m3)1.231.23Tensile Strength (MPa)2652.8Bed temperature(oC)60-7060-70Glass transition temperature(oC)8060Component design often involves managing information about materials with varying properties. One method to address this challenge is utilizing a Voxlizer-based slicing approach, enabling the manipulation of material distribution based on their respective volumes. Figure 3 illustrates the gradient interface within the investigated FGMs. Specifically, Figure 3a demonstrates a comparison between a conventional two-nozzle Multi-material AM process and the Multi-material FGM fabrication technique.

Specifically, Figure 3a demonstrates a comparison between a conventional two-nozzle Multi-material AM process and the Multi-material FGM fabrication technique. This comparison highlights the distinctive capabilities of the Multi- material FGM approach in creating graded interfaces within the fabricated structures. In Figure 3b, various combinations of different materials within the FGM are displayed. The Multi-material FGM is implemented in this study using the MEX technique.72Figure 3: Gradient transition of FGM interfaces, (a) conventional versus FGM material combination, (b) components with different material combinations [31]Specimen Preparation and Fabrication The fatigue test specimen in this study follows the guidelines outlined in ASTM standard E606M [32], as depicted in Figure 4. The specimen dimensions, including length, width, thickness, and curvature radius, are 135 mm, 20 mm, 5 mm, and 10 mm, respectively.

The specimen dimensions, including length, width, thickness, and curvature radius, are 135 mm, 20 mm, 5 mm, and 10 mm, respectively. Figure 5 visually represents the 3D printed FGM specimen.Figure 4: Dimension of the specimen as per E606MFigure 5: 3D printed FGM specimenExperimental setup73In this study, uniaxial tensile and fatigue tests were conducted on plastics using a Testresources 810E4 [33] load frame equipped with a 15 KN load cell, following the test procedures outlined in the ASTM E606M standard. The testing setup involved a closed-loop servo-hydraulic machine of the specified model. The Newton Testware [34] interface was utilized to control and input all testing variables, such as frequency, amplitude, and average load. The testing device consisted of two grasping heads, with the upper head serving as a simple grasping mechanism and the lower head responsible for applying the load to the specimen [35].

The testing device consisted of two grasping heads, with the upper head serving as a simple grasping mechanism and the lower head responsible for applying the load to the specimen [35]. A PID controller was employed to maintain the desired stress level during the tests. The experimental test setup is depicted in Figure 6.Figure 6: Test setupTensile testAn experimental investigation was conducted to gain a deeper understanding of the behavior of additively manufactured composite materials when subjected to different structural stresses. The study involved the use of tensile test samples, which were subjected to uniaxial loading until they reached the point of catastrophic failure within the gauge length area. The UTS values of the specimens were recorded and compared for different materials, including PLA, FGM, and TPU, all of which were printed using fixed processing parameters. The UTS values obtained for these materials are presented in Table 3.

The UTS values obtained for these materials are presented in Table 3. Furthermore, Figure 7 displays load-distance diagrams for PLA, TPU, and FGM specimens. Observations from the experiments indicate that the strength of the FGM material is lower than that of PLA but higher than that of TPU. These findings highlight that utilizing FGM can be an effective means of enhancing the strength of weak MEX parts.Table 3: UTS (MPa) of 3D printed samplesSample No. 1 2 3 AveragePLA 47.46 47.63 47.31 47.46FGM 30.88 30.70 31.20 30.92TPU 12.50 12.16 12.64 12.43742500TPU2000PLAFGM)N1500( d a o L1000500000.511.5Displacement (mm)Figure 7: Load vs DisplacementFatigue testTo assess the fatigue life of FGMs under specific conditions, a tension-based fatigue test was performed. The test involved varying parameters such as stress levels, stress ratio, and frequency.

The test involved varying parameters such as stress levels, stress ratio, and frequency. Stress levels of 80%, 60%, 40%, and 20% of the UTS were considered, with a stress ratio of 0.1. The frequency used in the test was 3 Hz. To ensure the consistency and reliability of the results, three case levels were examined, where FGM specimens were subjected to different stress levels. The test results were plotted using S-N curves. Moreover, curves are fitted with logarithmic trend line as depicted in Figure 8. The S-N curves exhibited slight variations, suggesting the presence of imperceptible anomalies such as voids, cracks, and porosity that may occur during the FGM deposition process. Notably, the specimens tested at 20% of UTS demonstrated a longer fatigue life, and the fatigue life of all samples was measured up to 140,000 cycles. Furthermore, it was observed that the specimens tested at 80% and 60% of UTS displayed characteristic behavior of low-cycle fatigue (LCF).

Furthermore, it was observed that the specimens tested at 80% and 60% of UTS displayed characteristic behavior of low-cycle fatigue (LCF). Conversely, the specimens tested at 40% and 20% of UTS exhibited behavior associated with high-cycle fatigue (HCF). The fatigue life of the FGM specimens was found to be more sensitive to the applied load levels. These research findings hold significant importance as they provide a comprehensive analysis of how stress levels influence the material characteristics of FGM parts. Moreover, the utilization of multiple case levels ensures the reliability and repeatability of the experimental outcomes, thereby enhancing the validity and robustness of the findings.7525) a P M2015Case I Case II Case III Log. (Case I) Log. (Case II) Log.

(Case I) Log. (Case II) Log. (Case III)(s s e r t S1050020000400006000080000100000120000140000160000Fatigue life (Cycles)Figure 8: S-N curve for experimental data with case I, II and IIIThe Statistical Methods of S-N Curve ModelsThe S-N curve plays a fundamental role in the estimation of material fatigue life. Various models have been developed to characterize experimental fatigue data subjected to constant amplitude loading and stress ratio. Researchers have employed a range of statistical models, including Basquin's Model, Weibull model [36], Sendeckyj[37] , Kim and Zhang model [13], and ANN [38], for the prediction of fatigue life.

Researchers have employed a range of statistical models, including Basquin's Model, Weibull model [36], Sendeckyj[37] , Kim and Zhang model [13], and ANN [38], for the prediction of fatigue life. These models aim to capture the relationship between stress amplitude and the number of cycles to failure, enabling accurate fatigue life predictions for different materials and loading conditions.Basquin’s Model It is widely employed by researchers for the prediction of fatigue life in additively manufactured components due to its simplicity and applicability to both polymers and composite materials. In a study conducted by Andrzejewska et al. [18] fitted the S-N curve for PLA using Basquin's model for specimens fabricated through MEX and injection molding processes.

In a study conducted by Andrzejewska et al. [18] fitted the S-N curve for PLA using Basquin's model for specimens fabricated through MEX and injection molding processes. Basquin's model is mathematically represented by equation 1, where 𝜎a denotes stress amplitude, C represents the fatigue strength coefficient, b corresponds to the fatigue strength exponent, and Nf represents the number of cycles to failure. This model describes the linear relationship between the applied stress cycles and the number of cycles to failure, presented on a logarithmic scale. Basquin's model can be viewed as a power law relationship capturing the fatigue behavior of the material under consideration.𝑏 = 𝐶𝛥𝜎 𝑁𝑓(1)In the Basquin’s equation, Δσ represents the applied peak stress, while C and b are the model fitting parameters or constants. The parameter Nf corresponds to the number of cycles at failure. When plotted on logarithmic-logarithmic coordinates, the Basquin’s equation yields a straight line.

The parameter Nf corresponds to the number of cycles at failure. When plotted on logarithmic-logarithmic coordinates, the Basquin’s equation yields a straight line. Equation (2), which is commonly employed for fatigue prediction, can be expressed as a linear logarithmic equation, as shown below:log Δσ + b log Nf = log C(2)76Results and DiscussionFollowing the successful fabrication of novel FGM specimens through the MEX process, a comprehensive evaluation of their performance was conducted. This evaluation encompassed various mechanical tests, including uniaxial tensile testing to assess their strength under axial loading and fatigue behavior analysis under the process parameters. Additionally, a statistical model based on Basquin's model was employed to predict the fatigue life of the FGM specimens. Following the acquisition of experimental data, the fatigue life of the specimens was determined by utilizing Basquin's law equation.

Following the acquisition of experimental data, the fatigue life of the specimens was determined by utilizing Basquin's law equation. The values of C and b in the Basquin's model can be obtained through a process of fitting the experimental fatigue life data to the mathematical equation 1. This process involves statistical analysis and regression techniques to determine the optimal values of C and b that best represent the relationship between stress level and fatigue life are represented in Table 4.Table 4: Values of C and b of Basquin’s relation and R2 correlationSamplescase I case II case IIIC128.56 126.74 134.75b0.1722 0.1738 0.1786Experimental (R2) 0.9947 0.9978 0.9973Basquin’s Model (R2) 0.9549 0.9608 0.9608% Difference 4 3.7 3.65The specimens were tested at stress levels corresponding to 80%, 60%, 40%, and 20% of the nominal UTS.

The calculated fatigue life based on the experimental data was compared with the predictions obtained from the Basquin's model, and the S-N curve was plotted using a logarithmic numerical function as shown in Figure 9. In Case I, the parameters C and b in the Basquin's model were determined to be 128.56 and 0.1722, respectively as presented in Table 4. Notably, the specimens tested at the 20% UTS stress level exhibited the highest fatigue life, exceeding 140,000 cycles. However, according to Basquin’s model predictions, the same stress level should yield a significantly higher number of cycles, specifically 1,727,342 cycles. The coefficient of determination (R2) was used to assess the statistical variation between the dependent and independent variables. The R2 value for the experimental fatigue life was determined to be 0.9947, indicating a strong correlation.

The R2 value for the experimental fatigue life was determined to be 0.9947, indicating a strong correlation. Similarly, the R2 value for the predicted fatigue life based on the Basquin's model was 0.9549, suggesting a good line of agreement with experimental. The percentage error between the experimental and predicted fatigue life cycles was determined to be 4% and curves are fitted with logarithmic trend line. Eventually, in Case I, it was observed that the experimental fatigue life and the predicted fatigue life cycles based on the Basquin's model exhibited some similarity. These findings hold scholarly significance as they provide insights into the comparison between experimental data and the predictions obtained from a well-known fatigue life prediction model, further contributing to the understanding of the fatigue behavior of the specimens.7725) a P M2015y = -2.803ln(x) + 38.925 R² = 0.9947Experimental Data Basquin's Model Log. (Experimental Data) Log.

(Experimental Data) Log. (Basquin's Model)(s s e r t S10y = -1.933ln(x) + 32.21 R² = 0.9549501000000 Fatigue Life (Cycles) Figure 9: S-N curve for Experimental data and Basquin’s Predicted model for case I050000015000002000000In Case II, the parameters C and b in the Basquin's model were determined to be 126.74 and 0.1738, respectively, as presented in Table 4. The fatigue life of the specimens tested at the 20% UTS stress level was found to be 140,000 cycles through experimental testing. Conversely, the predictions from the Basquin's model indicated a fatigue life of 1,394,213 cycles, as demonstrated by the S-N curve in Figure 10. For Case II, the R2 value for the experimental fatigue life was determined to be 0.9978, indicating a high degree of correlation. Similarly, the R2 value for the predicted fatigue life based on the Basquin's model was 0.960849, suggesting a reasonably within the expected region of the variation.

Similarly, the R2 value for the predicted fatigue life based on the Basquin's model was 0.960849, suggesting a reasonably within the expected region of the variation. The percentage error between the experimental and predicted fatigue life cycles was found to be 3.7% and curves are fitted with logarithmic trend line.25Experimental Data20Basquin's Model) a P M15y = -2.815ln(x) + 39.004 R² = 0.9978Log. (Experimental Data)Log. (Basquin's Model)(s s e r t S10y = -1.983ln(x) + 32.534 R² = 0.960850050000010000001500000Fatigue Life (Cycles)Figure 10: S-N curve for Experimental data and Basquin’s Predicted model for case II78In Case III, the Basquin's model parameters were calculated to be C = 134.75 and b = 0.1786, as presented in Table 4. The experimental testing revealed a fatigue life of 140,000 cycles for the specimens subjected to a stress level of 20% UTS.

The experimental testing revealed a fatigue life of 140,000 cycles for the specimens subjected to a stress level of 20% UTS. However, Basquin’s model predicted a fatigue life of 1,315,760 cycles, Moreover, curves are fitted with logarithmic trend line as depicted in Figure 11. For Case III, the R2 value for the experimental fatigue life was found to be 0.9973, indicating a strong correlation between the measured and observed data. Similarly, the R2 value for the predicted fatigue life based on the Basquin's model was 0.9608, suggesting a reasonably consistent with the experimental data. The percentage error between the experimental and predicted fatigue life cycles was calculated to be 3.5%.25Experimental Data20Basquin's Model) a P M15y = -2.887ln(x) + 39.848 R² = 0.9973Log. (Experimental Data)Log.

(Experimental Data)Log. (Basquin's Model)(s s e r t S10y = -2.037ln(x) + 33.191 R² = 0.9608500200000400000600000800000100000012000001400000Fatigue life (Cycles)Figure 11: S-N curve for Experimental data and Basquin’s Predicted model for Case IIIComparing Case I, Case II, and Case III, all three cases demonstrated high R2 values, indicating significant correlations between the experimental and predicted fatigue life. Furthermore, the percentage errors for all cases were relatively low, with Case I having a percentage error of 4%, Case II having a percentage error of 3.7%, and Case III having a percentage error of 3.5%. These findings illustrate the effectiveness of the Basquin's model, with the calculated values of C and b, in predicting the fatigue life of the FGM specimens across different stress levels. In summary, all three cases exhibited strong correlations between the experimental and predicted fatigue life.

In summary, all three cases exhibited strong correlations between the experimental and predicted fatigue life. The Basquin's model, with the calculated parameters, accurately predicted the fatigue behavior of the FGM specimens tested at various stress levels. These findings contribute to the understanding of fatigue characteristics in FGM materials and highlight the reliability of the Basquin's model in estimating fatigue life under different stress conditions.Validating Basquin's Model for Fatigue Life Prediction in Additively Manufactured Polymeric FGM MaterialsTo validate the applicability of the Basquin's model to FGM materials, a comprehensive study was conducted involving three different stress levels: 90%, 70%, and 50% of the UTS. The research approach involved initially predicting the fatigue life using the Basquin's model and estimating the corresponding values of C and b. Subsequently, experimental testing was conducted to obtain the79actual fatigue life data.

Subsequently, experimental testing was conducted to obtain the79actual fatigue life data. The first step in the validation process was to apply the Basquin's model to predict the fatigue life of the FGM specimens at aforementioned stress levels. Based on this initial prediction, the values of C and b were calculated for each stress level. These values served as crucial parameters for further analysis. Next, experimental fatigue tests were conducted on the FGM specimens under the specified stress levels. The obtained experimental data was then compared to the predictions made by Basquin’s model. To quantify the correlation between the experimental and predicted fatigue lives, a power function was utilized to plot the S-N curve and fit it to the experimental data.

To quantify the correlation between the experimental and predicted fatigue lives, a power function was utilized to plot the S-N curve and fit it to the experimental data. The R2 was then computed, Moreover, curves are fitted with power trend line as depicted in Figure 12.30Baquin's25R² = 0.9992Experimental20Power (Baquin's )) a P M15Power (Experimental)(s s e r t S10R² = 0.975150010002000300040005000600070008000Fatigue life (Cycles)Figure 12: Comparison of Basquin’s model with experimental dataThe results of the validation study demonstrated a remarkable agreement between Basquin’s model predictions and the experimental fatigue life data. The R2 value for the Basquin's model was calculated to be 0.9991, indicating a strong correlation between the predicted and experimental fatigue lives. This high R2 value demonstrates the excellent agreement between the model's predictions and the actual experimental data.

This high R2 value demonstrates the excellent agreement between the model's predictions and the actual experimental data. Similarly, the experimental R2 value was determined to be 0.9763, providing further evidence of the consistent relationship between the experimental results and the predicted values. These findings, as presented in Table 5, confirm the reliability and accuracy of the Basquin's model in predicting fatigue life.Table 5: Values of C and b of Basquin’s relation and R2 correlationC129.2b0.1749Experimental (R2)0.9763Basquin’s Model (R2)0.9953% Difference 1.90%The small percentage error difference of 1.9% between the Basquin's model predictions and the experimental results confirmed the suitability and accuracy of the model for predicting the fatigue life of FGM materials subjected to different stress levels.

These findings provide strong evidence that the Basquin's model is a reliable and effective tool for assessing the fatigue behavior of FGM80materials. In conclusion, the validation study demonstrated that the Basquin's model, with the calculated values of C and b, provided an excellent fit for predicting the fatigue life of FGM specimens under various stress levels. The high R2 values and low percentage error difference underscores the robustness and accuracy of the model, validating its suitability for assessing the fatigue performance of FGM materials in engineering applications.ConclusionsThis article presents a concise overview of an experimental study conducted on FGM specimens manufactured using the MEX process. The study focuses on the characterization of tensile properties and fatigue behavior, with particular emphasis on predicting fatigue life.

The main findings of this research can be summarized as follows:FGMs exhibited promising mechanical properties in terms of tensile strength and fatigue behavior.FGMs exhibited intermediate strength between PLA and TPU, suggesting their potential for strengthening weak MEX components.Fatigue testing at different stress levels demonstrated that FGM specimens exhibited both low-cycle fatigue and high-cycle fatigue behavior.The fatigue life of FGM specimens was more sensitive to load levels, with longer fatigue life observed at lower stress levels.Basquin's model demonstrated strong correlation and reliability in predicting fatigue life for FGM.The comprehensive analysis of stress levels' influence on FGM material characteristics provides valuable insights for designing and optimizing FGM parts in various applications.References[1][2]S.

43, no. 12, pp. 3027–3036, Dec. 2020, doi: 10.1111/FFE.13324.[36][37] A. K. Vanhari, E. Fagan, and J. Goggins, “A novel estimation method for fitting fatigue data in the composite wearout model,” Compos Struct, vol. 287, p. 115384, May 2022, doi: 10.1016/J.COMPSTRUCT.2022.115384.[38] H. Si, Z. Zhang, O. Huseynov, I. Fidan, S. R. Hasan, and M. Mahmoud, “Machine Learning-Based Investigation of the 3D Printer Cooling Effect on Print Quality in Fused Filament Fabrication: A Cybersecurity Perspective,” Inventions 2023, Vol. 8, Page 24, vol. 8, no. 1, p. 24, Jan. 2023, doi: 10.3390/INVENTIONS8010024.83Statements & DeclarationsFunding: This research project is based upon work supported by the Center for Manufacturing Research (CMR) and Mechanical Engineering Department. This support is greatly appreciated by the authors.Author Contributions: All authors contributed to various portions of the research project, mostly, however, by Ismail Fidan.

This support is greatly appreciated by the authors.Author Contributions: All authors contributed to various portions of the research project, mostly, however, by Ismail Fidan. Methodology, experimentation, and analysis were mostly done by Suhas Alkunte and Mithila Rajeshirke. The initial draft was written by Suhas Alkunte and Mithila Rajeshirke; revised and finalized by Ismail Fidan. All authors read and approved the final manuscript.Conflicts of Interest: The authors declare no relevant financial or non-financial interests.84.

tigue performance enhancement of selectively laser melted aluminium alloy by heat treatmentI. Maskery,1, ∗ N.T. Aboulkhair,2 C. Tuck,1 R.D. Wildman,1 I.A. Ashcroft,1 N.M. Everitt,2 and R.J.M. Hague1 1Manufacturing and Process Technologies Research Division, Faculty of Engineering, University of Nottingham, Nottingham NG7 2RD, UK 2Materials, Mechanics and Structures Research Division, Faculty of Engineering, University of Nottingham, Nottingham NG7 2RD, UKWe measured the stress-strain behaviour and fatigue performance of the aluminium alloy Al-Si10-Mg manufactured by selective laser melting (SLM). This process, specif- ically the rapid cooling of the metal from its molten state, results in a ﬁne microstruc- ture, generally providing high hardness but poor ductility. We used a heat treatment to alter the microstructure of the material from its as-built state. This signiﬁcantly improved the ductility and fatigue performance.

We used a heat treatment to alter the microstructure of the material from its as-built state. This signiﬁcantly improved the ductility and fatigue performance. The elongation at break for the heat treated material was nearly three times greater than that observed for the as-built material, and the fatigue strength at 106 cycles was around 1.6 times as high. Com- bined with the design freedoms of additive manufacture, this development increases the suitability of lightweight SLM parts for use in the aerospace and automotive sectors, where good fatigue performance is essential.I.INTRODUCTIONSelective laser melting (SLM) is an additive manufacturing (AM) process that constructs parts through the repeated deposition and melting of metal powder layers. Because of this layerwise construction, SLM allows parts with very complex geometries to be made, including those that come from topology optimisation (TO) algorithms[1–4] or include lattice structures of tessellating cells[5–13].

Complex, weight minimising geometries can be manufactured with no additional expense, and often considerably reduced processing time, than more traditional forms, i.e. those designed for conventional processes such as casting or milling. Lattices also have the potential to provide high levels of energy absorption under static and dynamic loading and, compared to TO methods, may oﬀer more robust solutions to problems which include uncertainty in the loading conditions or have multiple objectives. These features make SLM parts particularly well suited to the automotive and aerospace sectors, where the minimisation of component weight is a strong and persistent requirement.For TO and lattice structures to be used eﬀectively, and for the SLM process to be adopted into the relevant manufacturing sectors, the properties of SLM materials must be well understood.

Authors have examined many properties of SLM materials[14–21], but there exists only a handful of investigations into their fatigue performance[22–27]. Of these, only Siddique et al.[22] and Brandl et al.[25] have investigated the fatigue performance of SLM Al-Si alloys, and only Brandl et al. have examined Al-Si10-Mg in particular.∗email: ian.maskery@nottingham.ac.uk1017Siddique et al. investigated Al-Si12, focussing on the eﬀects of laser energy density, base plate temperature, and a post-build stress relief on the mechanical properties. Brandl et al. examined Al-Si10-Mg, likewise looking at the eﬀects of base plate temperature and a post-build heat treatment, in their case a T6 treatment which is often used to harden and strengthen conventionally produced Al-Si alloys. A few key ﬁndings of Siddique et al. and Brandl et al. are provided in section III of this work, where they are discussed in relation to our results.

A few key ﬁndings of Siddique et al. and Brandl et al. are provided in section III of this work, where they are discussed in relation to our results. It is worth noting that in both of these previous investigations the SLM fatigue specimens were subject to mechanical ﬁnishing to reduce their surface roughness prior to testing. To provide results more representative of as-manufactured SLM parts, our specimens were tested with their relatively rough surfaces intact.Heat treatments have been shown to signiﬁcantly alter the microstructure and hardness of SLM Al-Si10-Mg[16]. This material, with an inherently ﬁne microstructure arising from the SLM manufacturing process, responds diﬀerently to T6 or ‘precipitation hardening’ In the as-built condition, the heat treatments than when it is conventionally processed. microstructure is characterised by ﬁne α-Al grains with Si continuously segregated at the grain boundaries.

microstructure is characterised by ﬁne α-Al grains with Si continuously segregated at the grain boundaries. This material has a hardness of 109.9 ± 0.9 HV[16], which is high compared to the 86 HV of conventional cast and aged Al-Si alloy[14], but it is also relatively brittle (as seen in the stress-strain behaviour in section III). Enhancing the ductility of the material, by coarsening its microstructure with a post-manufacture heat treatment, could potentially increase the fatigue strength.The main aim of this investigation is to determine how a T6-like heat treatment aﬀects the microstructure, stress-strain behaviour and fatigue performance of SLM Al-Si10-Mg. Determining the relationship between these properties will add to the understanding of the material, and will support the development of eﬀective post-manufacture processes for improving the mechanical performance of SLM alloys.II. EXPERIMENTAL DETAILSA.

EXPERIMENTAL DETAILSA. Selective laser melting of test specimensA series of Al-Si10-Mg tensile test specimens and fatigue specimens were produced using a Renishaw AM250 SLM machine. The tensile tests were conducted using dog bone specimens in accordance with ASTM standard E8/E8M[28]. Their gauge length and diameter were 45 mm and 9 mm, respectively. The fatigue specimens had circular cross section and their geometry corresponded to ASTM standard E466-07[29]; the dimensions are shown in ﬁgure 1 and a photograph of a manufactured specimen is provided in ﬁgure 2. Both types of specimen were manufactured with their long axis perpendicular to the plane of the SLM build platform.During SLM specimen production, the laser power was 200 W and the powder was de- posited in 25 µm layers prior to each laser scan. The build platform was held at 180 ◦C throughout the manufacturing process.

The build platform was held at 180 ◦C throughout the manufacturing process. The laser rastered across the parts with a hatch spacing of 130 µm, a point distance of 80 µm and an exposure time of 140 µs, providing an eﬀective scanning speed of 571 mm s−1. The hatch direction of each layer was rotated by 67◦ from the previous one. Focused at the powder bed, the laser had a circular spot of101836.81364040FIG. 1: Dimensions (in mm) of the fatigue specimens used in this work.support materialFIG. 2: SLM Al-Si10-Mg fatigue specimen. Note the additional 2 mm of supporting material at the specimen base, which was included to facilitate removal of the specimens from the SLM build platform.diameter (68 ± 2) µm.B. Heat treatmentThe SLM Al-Si1-Mg specimens were tested in two conditions; as-built and heat treated. The post-manufacture heat treatment comprised a solution treatment for 1 hour at 520 ◦C followed by a water quench and artiﬁcial ageing for 6 hours at 160 ◦C.

The post-manufacture heat treatment comprised a solution treatment for 1 hour at 520 ◦C followed by a water quench and artiﬁcial ageing for 6 hours at 160 ◦C. This treatment was found by Aboulkhair et al.[16] to modify the microstructure of SLM Al-Si10-Mg and reduce the hardness by nearly 30% from the as-built condition. This softening is in stark contrast to the eﬀect on conventionally produced material, where treatments of this kind result in hardening through the precipitation of β-Mg2Si particles.The eﬀect of the heat treatment on the SLM Al-Si10-Mg microstructure is shown in ﬁgure 3. The as-built material, with its characteristic melt pools and ﬁne microstructure50 m μSi particles(a) (b) as-built heat treated50 m μmelt pool interiormelt pool boundaryFIG. 3: Optical micrographs of SLM Al-Si10-Mg in its as-built (a) and heat treated (b) conditions.1019Strain(%)Tensilestress(MPa)2As-built2001404000100Heattreated3003FIG.

3: Optical micrographs of SLM Al-Si10-Mg in its as-built (a) and heat treated (b) conditions.1019Strain(%)Tensilestress(MPa)2As-built2001404000100Heattreated3003FIG. 4: Tensile stress-strain curves of as-built and heat treated SLM Al-Si10-Mg specimens.that results from the SLM process, is transformed into a coarser-grained material, where the Si has agglomerated into particles with mean diameter just over 1 µm.C. Mechanical testingUniaxial static tension tests were conducted using an Instron 5581 universal testing machine equipped with a 50 kN load cell. The tests were conducted according to ASTM standard E8/E8M[28]. The load was applied at a displacement rate of 8.33 × 10−3 mm s−1. A random white and black spatter pattern was applied to the tensile test specimens, allowing strain data to be collected using a video gauge.

A random white and black spatter pattern was applied to the tensile test specimens, allowing strain data to be collected using a video gauge. Three repeat tensile tests were made for the specimens in their as-built and heat treated conditions.Room temperature fatigue tests were conducted using an Instron 8801 servo-hydraulic fatigue testing machine with a 100 kN load cell. A stress-controlled testing mode was used, following ASTM standard E466-07[29]. The cycling stress was sinusoidal with a frequency of 30 Hz, and a range of maximum stress, Smax, from 63 MPa to 220 MPa. The R value, equal to the ratio of minimum to maximum applied alternating stress, was 0.1. The runout limit was 3 × 107 cycles. Three repeat fatigue tests provided the data at each stress level.III. RESULTS AND DISCUSSIONThe stress-strain curves from the static tension tests are presented in ﬁgure 4. There are two main noteworthy features.

RESULTS AND DISCUSSIONThe stress-strain curves from the static tension tests are presented in ﬁgure 4. There are two main noteworthy features. First, the ultimate strength of the material was reduced by (12 ± 3)% from 330 ± 10 MPa to 292 ± 4 MPa after heat treatment , which is consistent with previous results concerning the eﬀect of the heat treatment on the hardness[16] owing to the modiﬁed microstructure. Second, the ductility of the material was enhanced by the heat treatment. This is seen in the increase in strain-to-failure from (1.4 ± 0.3)% to (3.9 ± 0.5)%, an increase by a factor of 2.8 ± 0.7.10201.6Heattreated106105As-built51.87As-built-linearﬁt1008RunoutR=0.130010401.4107Heattreated-linearﬁtlog(Nf)log(Sa)20064108Nf(cycles)Maximumstress,Smax(MPa)23×107cyclesFIG. 5: S-N curves of as-built (black circles) and heat treated (red triangles) SLM Al-Si10-Mg specimens.

5: S-N curves of as-built (black circles) and heat treated (red triangles) SLM Al-Si10-Mg specimens. A log-log plot is provided inset, along with corresponding linear ﬁts.Figure 5 provides fatigue curves of the as-built and heat treated material. At every level of applied maximum stress, Smax, the heat treated material survived a greater number of cycles before failure. The heat treatment also yielded a runout specimen, surviving 3 × 107 cycles, at a stress level of 94 MPa, while as-built specimens at the same stress had lifetimes of between just 2 × 105 and 5 × 105 cycles. The heat treatment, therefore, which increased the ductility of the SLM material by a factor of nearly three, also signiﬁcantly improved its fatigue performance.Brandl et al.[25] performed similar fatigue tests on SLM Al-Si10-Mg. The as-built and heat treated specimens of Brandl et al. had fatigue strengths of around 122 MPa and 170 MPa, respectively, at 106 cycles.

The as-built and heat treated specimens of Brandl et al. had fatigue strengths of around 122 MPa and 170 MPa, respectively, at 106 cycles. Thus, those authors also reported enhanced fatigue strength brought about by heat treatment. The corresponding fatigue strengths in this work were lower, at 85 MPa and 134 MPa, respectively. However, in comparing these results, we must consider some important diﬀerences in the experimental procedures employed in the diﬀerent studies. The most notable diﬀerences relate to the heat treatment and surface roughness of the examined specimens. The heat treatment of Brandl et al. consisted of a solution treatment for 6 hours at 525 ◦C, followed by water quenching and artiﬁcial ageing for 7 hours at 165 ◦C. See section II B for the heat treatment used here; the solution treatment duration was much shorter, at just 1 hour.

See section II B for the heat treatment used here; the solution treatment duration was much shorter, at just 1 hour. According to Aboulkhair et al.[16], the solution treatment duration is likely to strongly aﬀect the microstructure and hardness of the material, therefore will be like to aﬀect the fatigue performance also. Secondly, the specimens of Brandl et al. were machined and polished1021As-built Heat treatedlog(A) 2.86 ± 0.092.6 ± 0.1B0.21 ± 0.02 -0.14 ± 0.02TABLE I: Fatigue curve ﬁtting parameters for as-built and heat treated specimens.in accordance with ASTM E466-07[29] (which speciﬁes a maximum surface roughness of 2 µm), whereas the specimens in this work retained their as-built surface roughness. The surface roughness of SLM Al-Si10-Mg has previously been reported in the range of 12 µm[30] to 24 µm[31].

The surface roughness of SLM Al-Si10-Mg has previously been reported in the range of 12 µm[30] to 24 µm[31]. Since stress concentrations at the specimen surface cause crack initiation during cyclic loading, the reduced surface roughness of Brandl et al.’s specimens is likely to result in better fatigue performance than exhibited by the specimens examined here.To provide more quantitative evidence regarding the fatigue performance of the as-built and heat treated material, numerical ﬁtting was performed on the S-N data. A relationship of the form proposed by Basquin[32] was used, which can be expressed as[33],Sa = A (Nf )B,where Sa is the stress amplitude, Nf is the number of cycles to failure, and A and B are constants. Taking base-ten logarithms provideslog(Sa) = log(A) + B log(Nf ).The constant B is then the gradient of the S-N data when plotted on a log-log basis. This is provided in the inset of ﬁgure 5, which also includes the ﬁt lines determined by linear regression.

This is provided in the inset of ﬁgure 5, which also includes the ﬁt lines determined by linear regression. The ﬁt parameters for the as-built and heat treated materials are provided in table I. The slope of the S-N log-log plot was reduced from -0.21 ± 0.02 to -0.14 ± 0.02, demonstrating the improvement in fatigue performance brought about by the heat treatment. Brandl et al. performed a diﬀerent form of statistical analysis on their SLM Al-Si10-Mg fatigue data, so a direct comparison can not be made in this regard between our work and theirs. The best source for comparison is the work of Siddique et al.[22], who examined SLM Al-Si12 and also ﬁtted their S-N fatigue data with a Basquin model. Their data yielded a log-log slope of -0.129, which agrees within error with that of our heat treated material. Agreement between our result and that of Siddique et al.

Their data yielded a log-log slope of -0.129, which agrees within error with that of our heat treated material. Agreement between our result and that of Siddique et al. was not expected, as the material composition and post-manufacture treatment were both diﬀerent (again, their specimens were machined to lower the surface roughness), but it is encouraging that similar materials manufactured by the same pro- cess behave comparably. Siddique et al. also observed a fatigue strength at 106 cycles of approximately 80 MPa, which is very close to the 85 MPa seen for the as-built material here.Other relevant results come from Buﬃ`ere et al.[34], who studied the related alloy Al- Si7-Mg0.3 made by casting. They found values of B from -0.168 to -0.131 for a range of Al-Si7-Mg0.3 alloys with diﬀering populations of pores. For the heat treated material studied here, the value of B is in the range obtained by Buﬃ`ere et al., indicating similarity of behaviour in our respective materials.

For the heat treated material studied here, the value of B is in the range obtained by Buﬃ`ere et al., indicating similarity of behaviour in our respective materials. This result is consistent with the similar chemical composition and microstructure of the two materials, i.e. an Al α phase interspersed with1022(1)(2)Si particles. We must be cautious in drawing conclusions from this comparison, however, since the manufacturing processes are quite diﬀerent, and the Si particles seen by Buﬃ`ere et al. were larger than those seen here (3.1 µm cf. ∼ 1 µm in this work) and were less well dispersed throughout the material.IV. CONCLUSIONSWe have shown that a heat treatment, of the kind which hardens and strengthens conventionally manufactured Al-Si alloys, results in reduced strength but signiﬁcantly increased ductility and fatigue strength in SLM material.

Brandl et al.[25] came to a similar conclusion regarding the eﬀect of a heat treatment on the fatigue performance, and presented fatigue strengths which were in excess of those seen here for the same alloy. However, our specimens were tested with unmachined surfaces, so our results are more rep- resentative of the properties readily achievable by SLM without mechanical post-processing operations. These can be time consuming even for relatively simple geometries, and may often be impossible for the kind of complex lightweight parts, i.e. topology-optimised structures and lattices, that SLM is capable of producing.

topology-optimised structures and lattices, that SLM is capable of producing. The results presented here, then, are novel and useful, and the information regarding the relationship of the microstructure and mechanical properties will aid the development of new heat treatments that max- imise the desirable material properties with the minimum expenditure of eﬀort and resources.To expand upon this work, we will examine the eﬀect of surface roughness of SLM parts on their fatigue performance. This will provide further insight into the respective importance of the microstructure and surface quality on the fatigue performance.AcknowledgmentsFunding was provided by the UK Technology Strategy Board. Thanks to Mark East,Mark Hardy, Joe White, Jason Greaves and Tom Buss.[1] Aremu, A. et al. The eﬀects of bidirectional evolutionary structural optimization parameters on an industrial designed component for additive manufacture. P. I. Mech. Eng. B- J. Eng.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFatigue Performance of Additively Manufactured Stainless Steel 316L for Nuclear Applications William Beard1,a, Robert Lancaster1,b, Jack Adams2,c, Dane Buller2,d 1Institute of Structural Materials, College of Engineering, Bay Campus, Swansea University, Swansea, SA1 8EN, United Kingdom 2Rolls-Royce Submarines, P.O. Box 2000, Raynesway, Derby, DE21 7XX, United Kingdom a790871@swansea.ac.uk , br.j.lancaster@swansea.ac.uk, cJack.Adams@rolls-royce.com, ddane.buller@rolls-royce.comKeywordsAdditive manufacturing; laser powder bed fusion; stainless steel 316L; low cycle fatigueAbstractAdditive manufacturing (AM) is a rapidly growing technology which is extending its influence into many industrial sectors such as aerospace, automotive and marine.

Recently the nuclear sector has considered AM in the production of nuclear reactor components due to its possible advantages over conventional manufacturing routes. This includes considerable cost savings due to less material wastage, the ability to produce complex near net shape components that conventional manufacturing processes are unable to achieve and a reduced manufacturing time. Initially, Stainless Steel 316L (SS316L) manufactured by laser powder bed fusion (LPBF) has been identified as a potential candidate. However, due to the transient nature of the microstructure it is now of fundamental importance to assess and understand the mechanical behaviour of the LPBF material.This paper will highlight some of the recent research at Swansea University in investigating the variation on the fatigue characteristics between wrought SS316L and LPBF processed SS316L material. This will include an extensive microstructural and fractographic investigation.

This will include an extensive microstructural and fractographic investigation. As LPBF material looks to replace conventionally manufactured equivalents, an understanding of how build integrity and orientation affects the mechanical properties of AM material is critical. Wrought and vertical LPBF material are to be assessed to understand how the microstructure controls the fatigue performance of LPBF SS316L material.21441. Introduction Additive manufacturing (AM) is a near net shape manufacturing process currently under consideration by the nuclear sector to replace conventional manufacturing techniques. The AM process offers many advantages in the production of parts compared to conventional manufacturing techniques and this has led to many industrial sectors such as medical, automotive and aerospace investing in the technology. Overall, AM has unrivalled design freedom and provides the ability to manufacture single or multiple components from a wide range of materials.

Overall, AM has unrivalled design freedom and provides the ability to manufacture single or multiple components from a wide range of materials. This allows the production of complex parts produced to near net shape, meaning there is no need for expensive tooling such as punches, dies or casting moulds. This therefore reduces the amount of processing steps which is not only seen to reduce manufacturing time but also improve component strength and integrity, as parts have fewer joining regions removing possible stress raising features [1]. However, challenges in understanding how the feedstock materials, processing variables and other contributing factors affect the mechanical properties and performance of a component are still required. The AM process typically builds a three-dimensional part on a layer by layer basis by utilising a high energy heat source to selectively melt and fuse two-dimensional slices of a component [2, 3].

The AM process typically builds a three-dimensional part on a layer by layer basis by utilising a high energy heat source to selectively melt and fuse two-dimensional slices of a component [2, 3]. The high energy heat source is controlled by digital data from computer aided design (CAD) to build each layer in a sequential manner. AM processes fall into two categories defined by the ASTM Standard F2792 [4]. The processes are characterised as directed energy deposition (DED) and powder bed fusion (PBF). This study focusses on laser powder bed fusion (LPBF) which utilises a laser as a thermal energy source that selectively fuses regions of a powder bed [2]. A schematic of the LPBF process can be seen in figure 1. LPBF is based on the complete melt and solidification of metallic powder to produce near fully dense components [5]. However, in most cases post processing in the form of heat treatments and machining is required.

However, in most cases post processing in the form of heat treatments and machining is required. Post process heat treatments such as hot isostatic pressing (HIP) [6, 7] can be used to reduce the number of defects such as porosity which can be catastrophic under mechanical loading.X•Y galvano meterF·theta lensGas InletInert atmosphereGas outletLoose powder Solid structuretMetal substratePowder platformFigure 1- Schematic of the LPBF Process [8]2145Many studies have been carried out on additive materials under static loads, however dynamic cyclic loading in the form of fatigue is not yet fully understood. These dynamic cyclic loading conditions are a requirement for many critical applications within the aerospace, marine and nuclear industry. Therefore, gaining an understanding of the fatigue behaviour of LPBF materials is of upmost importance.

Therefore, gaining an understanding of the fatigue behaviour of LPBF materials is of upmost importance. This paper aims to provide a further understanding into the fatigue characteristics of LPBF processed material in comparison with its wrought counterpart.2. Materials and Methods The material focussed on in this study is stainless steel 316L (SS316L). SS316L is an austenitic grade of stainless steel with a relatively high nickel content and moderate chromium content. SS316L is a marine grade corrosion resistant stainless steel with good weldability due to its low carbon content and moderate molybdenum content. These properties make it an ideal candidate to be manufactured by LPBF since it displays many similarities with welding.2.1. Material Variants and fabrication Two variants of SS316L were investigated in this study; a conventionally manufactured wrought variant and a LPBF variant.

Material Variants and fabrication Two variants of SS316L were investigated in this study; a conventionally manufactured wrought variant and a LPBF variant. The LPBF material was fabricated in the vertical (90°) build orientation using an EOS M290 machine using gas atomised (GA) SS316L powder with the nominal composition seen in table 1 and a powder particle size distribution of 5-60 μm. The processing parameters adopted to manufacture the samples were the standard parameters set by the machine manufacturer EOS to obtain samples with near full density. Four φ 20mm cylindrical samples were built in the vertical orientation to a height of 105mm. After building the cylinders were removed and hot isostatically pressed (HIPped) with the aim of further consolidating the material [6, 7].Table 1- Nominal composition of SS316L gas atomised powder (wt.%)Element Fe SS316L BalC 0-0.03Cr 16-18Mn 0-2Mo 2-3Ni 10-14P 0-0.045 0-0.03S2.2.

Microstructure Characterisation A Hitachi SU3500 scanning electron microscope (SEM) equipped with an electron backscatter diffraction (EBSD) unit was used to characterise the microstructural features of the two variants. EBSD scans were conducted with the SEM and EBSD settings displayed below in table 2. EBSD data was acquired using the tango plug-in included in the Channel 5 software. Careful noise reduction was performed prior to EBSD analysis whereby wild spikes and zero solutions were replaced. However, zero solutions were only conducted three times.

Careful noise reduction was performed prior to EBSD analysis whereby wild spikes and zero solutions were replaced. However, zero solutions were only conducted three times. Grain detection was performed using the grain area determination in tango, where no border grains were included and boundaries with a misorientation >10° were used for measurements of grain size.2146Si 0-1Table 2- SEM and EBSD settings used during EBSD data acquisitionSEM SettingsEBSD SettingsAccelerating Voltage Working Distance Step Size EBSD camera binning mode 4 x 4 Hough resolution Phases for acquisition Band detection mode20 KV 20 mm 1.5 μm60 Iron FCC, Iron BCC Optimised EBSD2.3. Strain Control Low Cycle Fatigue Testing Strain Controlled Low Cycle Fatigue (LCF) tests were performed on a servo-hydraulic mechanical test machine according to ASTM E606 [9] with specimens machined according to the dimensions depicted in figure 2.

After machining, the gauge sections of all specimens were circumferentially polished with emery papers of 320/800/1200 grit and 3 μm diamond suspension. Testing was conducted at 20°C under controlled laboratory conditions, and a 0.4s- 1 triangular waveform with an R ratio of -1 was employed under strain control. Maximum strain ((cid:304)MAX) values of 0.3 and 0.6% were used to investigate the effects of different loading conditions on the wrought and LPBF material. After testing was complete a Hitachi SU3500 SEM was used to inspect the fracture surfaces.Ml2 x l.75pFigure 2- Specimen drawing of LCF test pieces with dimensions in mm2.4. Surface Roughness Measurements Prior to testing, an Alicona Infinite Focus microscope was used to measure the roughness topography and the related parameters of all samples according to ISO 4287; ISO 4288.

Material Characterisation Figure 3 presents the EBSD micrographs for both the wrought and vertical LPBF variants in both the XY and XZ orientations. The XZ Plane is orientated in the build direction. Overall both the wrought and vertical LPBF material shows an austenitic Face Centre Cubic (FCC) microstructure with minute amounts of retained Body Centre Cubic (BCC) (cid:303)-ferrite distributed throughout and this is shown clearly in figure 4. Grain measurements, kernel average misorientation (KAM), quantity of (cid:520)(cid:22)(cid:3)(cid:69)(cid:82)(cid:88)(cid:81)(cid:71)(cid:68)(cid:85)(cid:76)(cid:72)(cid:86) (twinning boundaries) and amount of BCC ferrite phase for the wrought and vertical samples are reported in table 4. From the table it is clear that there is a higher percentage of BCC ferrite in the wrought sample.

From the table it is clear that there is a higher percentage of BCC ferrite in the wrought sample. Figure 4a shows that the BCC ferrite phase has a needle like morphology in the wrought microstructure whereas the LPBF material, in figure 4b and c, shows very small amounts of randomly orientated BCC ferrite phase along the grain boundaries. Furthermore, figure 4 shows substructure as highlighted by the silver colour indicator. In the wrought material, very little substructure can be observed whereas the LPBF material shows large amount of substructure within the grains.The wrought microstructure seen in figure 3a shows an equiaxed grain morphology where mechanical properties would be expected to be predominantly isotropic. Whereas the LPBF material in figure 3b shows a strongly anisotropic structure as the XY face shows an elongated microstructure due to epitaxial grain growth along the build direction and an equiaxed grain morphology perpendicular to the build orientation.

This columnar or epitaxial grain growth has most likely arisen due to the complex thermal history occurring in the LPBF process [3]. Since LPBF is a layer by layer manufacturing process, a single layer of metal is cast upon a previous single layer. This causes time dependant temperature profiles and often results in an alloy experiencing repeated solid state and liquid solid phase transformations [3]. Additionally LPBF is a relatively rapid solidification process with reported cooling rates of 104 K/s and heat flow in the build direction frequently results in columnar microstructures [3].2148The microstructure displayed in figure 3b is after the LPBF material has been HIPped. During the HIP cycle it is likely that recrystallisation has not occurred due to the lower temperature used in comparison to previous studies [5, 10] and therefore no microstructural transformation has occurred.

However, the significant local deformation has been successful in closing up any small pores to give a fully dense structure.zX J-vFigure 3- EBSD microstructure in XY and Z direction a) Wrought b) VerticalFigure 4- EBSD Phase characterisation of a) wrought b) Vertical LPBF Z c) Vertical LPBF XY where blue shows FCC austenite structure and red shows BCC ferrite structure. Additional sub-grain boundaries shown as silver.Table 4 highlights the topological parameters recorded in each variant. When considering the aspect ratio of the differing grain structures the wrought sample displays a value closest to 1, indicating the most equiaxed structure, with the LPBF vertical XY face displaying the lowest value. This is most likely attributed to the thermal gradients occurring during the LPBF process. It is important to note that significant differences are seen between the wrought and LPBF material when looking at average KAM and (cid:520)(cid:22).

It is important to note that significant differences are seen between the wrought and LPBF material when looking at average KAM and (cid:520)(cid:22). In terms of average KAM the wrought material shows a much lower average KAM value in comparison to the LPBF material indicating a higher residual stress in the LPBF material. In terms of the (cid:520)(cid:22) value the wrought build appears to show a much larger amount of twinning boundaries in comparison to the LPBF material. This is to be expected as twinning grain boundaries are often seen in wrought SS316L [11] however previous studies have highlighted twinning induced plasticity occurring in LPBF SS316L [12].2149Table 4- Grain measurementsSampleWrought Vertical XZ Vertical XYAverage Grain Size ((cid:541)(cid:80)2) 1065.5 408.51 646.19Number of Grains406 1143 848Aspect Ratio0.72 0.59 0.38Average KAM (°)0.15 0.65 0.55(cid:520)(cid:22)(cid:3) (%)49.0 0.573 0.531Amount of BCC phase (%) 1.5 0.2 0.43.2.

Surface Roughness Surface roughness measurements can be seen in table 5. Previous research [13] has identified strong correlations of fatigue life with surface roughness, however no direct trend was observed in the data collected here. However, it is important to note that the circumferential polish is more likely to affect fatigue performance since the polish marks align with the crack growth direction perpendicular to the applied load [14].Table 5- Surface Roughness Measurements on SamplesHeight ParametersAmplitude Distribution ParametersSampleW-2 W-3 V-2 V-4Rv Ra (μm) (μm) 0.87 7.11 4.04 3.07 5.36 1.14 0.05 1.08 8.26 3.73 4.53 6.25 1.38 -0.11 0.99 7.04 3.91 3.13 5.71 1.25 0.07 -0.55 0.83 6.20 2.49 3.72 4.84 1.04Rt (μm)Rp (μm)Rz (μm)Rq (μm)Rsk (μm)Rpk (μm) 1.33 1.45 1.24 0.79Mr2 Rvk (μm) (%) 1.22 2.56 12.07 87.04 1.51 3.01 16.89 88.53 1.07 3.10 11.95 89.32 87.37 1.20 2.60 8.1Rk (μm)Mr1 (%)3.3.

Strain Control Low Cycle Fatigue (LCF) TestingFigure 5 presents the stress evolution with time for the eight LPBF specimens tested in this research. In both material variants, it can be seen that there is initial cyclic softening occurring in the first 1000 cycles. Over time the stresses start to level and become more stabilised until final fracture occurs, as indicated by the drop in maximum stress.

: 11: - ~ • II _,;;-•••I.A------------W-4 V-4 , .. ,, ,, ,, ,, ..,14005000500010000150002000025000NFigure 5- Maximum and minimum stress evolution behaviour for (cid:304)MAX = 0.3% and 0.6% testsThe characteristic hysteresis strain-stress loops for (cid:304)MAX = 0.3% are shown in figure 6. The monotonic loops (N=1) highlight both the Youngs’ modulus (E) of the material and the yield stress ((cid:305)ys) behaviour. Both variants appear to have a similar modulus of 200MPa however differing yield responses are seen between the wrought and vertical LPBF material. The wrought material is seen to yield at a lower value of around 200MPa in comparison to the vertical LPBF material which yields around 280MPa. The stabilised (½ Nf) loops highlight the stabilised stress range (Δ(cid:305)STAB) of the materials and show further evidence of cyclic softening occurring in the material.

The stabilised (½ Nf) loops highlight the stabilised stress range (Δ(cid:305)STAB) of the materials and show further evidence of cyclic softening occurring in the material. Additionally, the Δ(cid:305)STAB values show that the (cid:305)ys in the wrought material has led to a reduced Δ(cid:305)STAB in comparison to the vertical LPBF material.The characteristic hysteresis loops corresponding to (cid:304)MAX = 0.6% are shown in figure 7. This shows a similar response to the 0.3% regime whereby all materials show a similar E with the wrought material yielding at a lower value. This yielding has also led to a reduced Δ(cid:305)STAB for the wrought material.Fatigue cracks can nucleate early in LCF due to the large plastic strains occurring in localised regions [15]. The total strain amplitude can be resolved into plastic and elastic strain components. Generally speaking, the width of the hysteresis loops show the amount of plasticity at the given test strain amplitude.

Generally speaking, the width of the hysteresis loops show the amount of plasticity at the given test strain amplitude. As such, in both the (cid:304)MAX = 0.3% and 0.6% regime the wrought material variants are seen to show a wider hysteresis loop than the vertical samples indicating a greater amount of plasticity in these variants. In addition, the results obtained were lower than expected and this is most likely explained by the circumferential polished surface finish [14].

Fractography The fracture surface of a metallic component subjected to LCF loading is typically comprised of three main areas [15] and these were identified in all fatigue fracture surfaces analysed by SEM microscopy in figure 8. These include a period of crack nucleation, followed by a zone of crack propagation and the final failure region. The origin of fatigue damage for all samples was related to surface features most likely due to surface roughness and the circumferential polish on the samples. A large area of crack growth can be observed in all fracture surfaces. However, as an R ratio of -1 was employed in this research, fatigue striations were not clearly visible at high magnification due to the fully reversed tensile and compressive forces applied.

However, as an R ratio of -1 was employed in this research, fatigue striations were not clearly visible at high magnification due to the fully reversed tensile and compressive forces applied. The final failure areas present a ductile morphology as evidenced by the presence of microscopic pores and necking on the outer surface of the specimens.2153(cid:959)(cid:2239)e0.30 0.33 0.36 0.35 0.37 0.36 0.44 0.44MlllhllliFigure 8- Fracture surfaces of fatigue samples a-b) Wrought 0.3% c-d) Vertical 0.3% e-f) Wrought 0.6% g-h) Vertical 0.64. Conclusion The LCF behaviour of wrought and vertical LPBF SS316L was assessed and results showed that the vertical LPBF material performed equally to its wrought equivalent.

Conclusion The LCF behaviour of wrought and vertical LPBF SS316L was assessed and results showed that the vertical LPBF material performed equally to its wrought equivalent. The initial microstructural characterisation revealed that the wrought material contains an isotropic equiaxed structure with a large amount of twinning boundaries whilst the LPBF material showed an anisotropic structure with epitaxial columnar grains orientated in the build direction. In addition, the LPBF material appeared to contain fewer twinning boundaries and a large amount of sub grain structure. Overall these features were found to have a limited influence on the fatigue properties of the material variants. This is most likely due to the circumferential polish of samples and the relative roughness that has been measured. However no direct correlation was found between roughness and fatigue life.

This is most likely due to the circumferential polish of samples and the relative roughness that has been measured. However no direct correlation was found between roughness and fatigue life. When closely looking at the hysteresis loops it was seen that the wrought material showed greater plasticity under loading than the vertical LPBF samples, and as such, would be expected to exhibit an inferior resistance to fatigue damage.5. Acknowledgements This research was funded by The Materials and Manufacturing Academy (M2A) supported by the European Social Fund through the Welsh Government. The provision of a research bursary, materials and supporting is gratefully acknowledged. Alicona surface roughness measurements were completed in the Welsh Centre for Printing and Coating (WCPC) and mechanical tests were performed at Swansea Materials Research and Testing Ltd. (SMaRT). information from Rolls-Royce plc2154References [1]Ford S, Despeisse M.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFatigue Properties of 3D Printed Maraging SteelNandhini Raju*, and David W. Rosen†Department of Mechanical and Aerospace Engineering, University of Central Florida, Orlando, FL 32816. †Digital Manufacturing and Design Centre, Singapore University of Technology and Design, Singapore 487372. †School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332.Corresponding author: nandhini@knights.ucf.eduAbstractThe objective of this paper is to investigate fatigue properties of maraging steel, in the fabricated, machined, and heat-treated conditions, printed by powder bed fusion. Samples were manufactured in an EOS M280 machine in the X and Z build directions. Manufactured samples were tested under four different conditions: fabricated, machined, heat-treated, and heat-treated and machined.

Manufactured samples were tested under four different conditions: fabricated, machined, heat-treated, and heat-treated and machined. Each condition was expected to have different fatigue properties. The maximum stress and number of cycles to failure results were compared to understand the influence of the different build orientations and conditions on fatigue properties and limits. Results showed that machining and heat treatment, individually and together, had significant effects on fatigue properties. Additionally, the selection of standards, selection of sample counts, and statistical analysis of results will be discussed along with the maraging steel fatigue properties.IntroductionAdditive manufacturing is also known as rapid prototyping or 3D printing. AM techniques have a significant influence over the automotive, aerospace, medical field. Also, it changed the total experience of the manufacturing itself.

AM techniques have a significant influence over the automotive, aerospace, medical field. Also, it changed the total experience of the manufacturing itself. AM has its unique set of advantages over conventional manufacturing such as less time for manufacturing, being capable to manufacture any complex designs, design freedom, and less material wastage. Due to these advantages, these techniques are getting much attention in industry and academia. Generally, AM techniques can be categorized into seven process categories[1]. Of interest to this paper, metal powder bed fusion (PBF) techniques are used commonly in many industries due to their high density and low porosity end products.

Of interest to this paper, metal powder bed fusion (PBF) techniques are used commonly in many industries due to their high density and low porosity end products. However, after melting, the material cools down rapidly, causing high-temperature gradients that leave residual stress in the material and influence the material’s strength especially fatigue strength.Materials such as maraging steels are used in aviation, motor, and military applications due to their high toughness, high ultimate tensile strength along with its high ductility, and weldability[2][3]. Several researchers are working on the mechanical and material properties to improve the strength of the printed steel. Fatigue properties are important for many applications. Generally, a fatigue test is done by applying cyclic loading which is less than the yield strength of the material at static loading. Cyclic loading can be tensile, compressive, or a combination of both.

Cyclic loading can be tensile, compressive, or a combination of both. Before material reaches the fracture stage, crack initiation and crack growth are developed in the831material[4]. Crack initiation can happen at the surface or subsurface region due to surface defects such as cracks, porosity, and inclusions. After initiation, a crack grows along the slip plane and is further propagated in the material. This can be predicted in microstructure analysis which can be indicated by striations on the surface [4]. When stress intensity is higher than the material’s strength it leads to fracture. Fatigue limit and endurance limit are determined after the collection of stress amplitude and life cycle data, which is typically plotted as the well-known S-N curve. Usually, materials used in an application are subject to loading conditions below the fatigue limit for safety considerations.

Usually, materials used in an application are subject to loading conditions below the fatigue limit for safety considerations. Usually conventionally manufactured maraging steels’ fatigue limit is known as 50% of its UTS (ultimate tensile strength) [5]. But the fatigue limit of maraging steel manufactured by additive manufacturing is 20-30% of its UTS which is much lower than conventionally engineering materials [6]. So, improving the fatigue limit for the additive manufacturing part is important especially for the materials used in harsh environments such as maraging steel.Maraging steels are known to be precipitation-hardened steels, with intermetallic precipitates forming as the result of aging heat treatments. These precipitates are formed from the various alloying elements, including Fe, Ti, Ni, Al, and Mo, with Ni3Ti predominating. The aging heat treatment used in this work was performed at 490 °C for 6 hours. Grade 300 maraging steel was used in this study.

The aging heat treatment used in this work was performed at 490 °C for 6 hours. Grade 300 maraging steel was used in this study. After aging, ultimate tensile strength increases from around 1100 MPa to 2000 MPa. Hardness improves by a similar amount, while elastic modulus increases by approximately 20 percent.Literature reviewFatigue testing can be performed with either bending loading or axial loading. The loading cycle can be symmetric, asymmetric, and random type according to stress and time. Usually, ferrous materials such as steel alloys have fatigue and endurance limits, that is finite lives, but non- ferrous alloys such as aluminum alloys may not have endurance limits. Usually, fatigue has three failure modes that consist of crack initiation, propagation, and catastrophic fracture. Generally, shear flow along the slip plane generates the intrusion and extrusion due to the loading cycle.

Generally, shear flow along the slip plane generates the intrusion and extrusion due to the loading cycle. These create surface defects and later it initiates the crack on the surface especially in sharp corners. Tensile loads can fasten crack propagation of the material and compressive stress closes the crack. So tensile load helps to open the crack and compress does the opposite. The repeated loading of tension and compression helps to propagate a crack in the material. Cross-sections of the fatigue failed material can be helpful to predict the propagations. Generally, fatigue tests are influenced by surface cracks, surface roughness, grain size, residual stress, and environment[4].Material such as maraging steel is high in hardness and it may produce surface cracks easily due to its brittleness. To reduce the number of surface defects, several attempts such as machining, shot peening, surface finishing were usually used.

To reduce the number of surface defects, several attempts such as machining, shot peening, surface finishing were usually used. Also, for conventionally engineered maraging steels, different aging treatments enhance the tensile strength, which leads to improved fatigue strength of the material[7]. Also, it is observed that decreasing the aging temperature of conventionally manufactured steels decreased the austenite fraction and interparticle spacing which resulted in improving strength. To improve the fatigue strength different attempts such as changing the chemical composition, nitriding, and shot-peening[8] the manufactured steel were done. Particularly, nitriding the conventional maraging steel increased the fatigue strength due to hardening[9]. As part of hardening, it formed compressive residual stresses in the surface, which832benefits fatigue strength.

As part of hardening, it formed compressive residual stresses in the surface, which832benefits fatigue strength. Different types of fatigue tests have been developed such as the ultrasonic fatigue test, very high cyclic test (VHCF) to understand the limitations of the materials [10]. From these tests, it is observed that inclusions were the origin of crack initiation [11]. Especially inclusions such as Ti-N and aluminate were the main reasons for the crack initiation. Also, even 0.6µm inclusions[12] lead to initial cracks in the surface and sub-surface region which lead to fracture.According to the authors’ knowledge, limited fatigue-related research is available for additively manufactured (AM) maraging steels. Suryawanshi et.al stated that fine cellular structure observed in SLM printed maraging steel showed higher strength when it was aged, and it was comparable with wrought material[2].

Suryawanshi et.al stated that fine cellular structure observed in SLM printed maraging steel showed higher strength when it was aged, and it was comparable with wrought material[2]. Intermetallic components (Fe-Mo based, Ni-based) present within the finer cellular structure were the main reason for the improved ultimate tensile strength, which eventually resulted in improving the fatigue limit of the material. Similar to aging, some researchers have noted that build orientation did affect the fatigue life[13], while other researchers reported that they observed no variation regarding build orientation in fatigue strength[5][14]. Due to the number of variables in PBF processes, using the same build with the same powder, same parameters, etc, was crucial to consolidate the results. Antunes et.al reported the propagation of crack growth in between the layer was significant. This proved the importance of bonding strength and usage of process parameters to manufacture the samples[15].

This proved the importance of bonding strength and usage of process parameters to manufacture the samples[15]. Sometimes process parameters played an important role to manufacture the quality dense parts. It was reported by Santos et.al that higher scan speeds lead to high porosity[16]. Little is known about the effects of build orientation and post-processing treatments on the fatigue behavior of the EOS MS1 maraging steel material. In this research, build orientation and different post-processing conditions such as- printed, printed and machined, heat-treated, and heat-treated and machined are considered in characterizing the fatigue life of the EOS MS1 material.Research MethodologyMaraging steel specimens were manufactured by the EOS M280 system using EOS MS1 powder at the Singapore University of Technology and Design. Four different post-processing conditions were examined including as-printed(A), printed and machined (B), heat-treated(C), and heat-treated and machined(D).

Four different post-processing conditions were examined including as-printed(A), printed and machined (B), heat-treated(C), and heat-treated and machined(D). According to the experimental campaign, samples for each condition were printed in the same build with the same powder to avoid the variance in the results.833Figure 1 flow chart of the proposed characterization method (reproduced with permission[17])After the selection of the different post-processing conditions, sample counts are selected based on the characterization method described in the flow chart in Fig.1, which was explained in detail in the authors’ previous work[17]. Sample counts were decided to provide a higher confidence interval based on the ASTM E122-17[18] standard and the selection of standards for AM was derived from the NISTIR 8005[19] technical report.

Sample counts were decided to provide a higher confidence interval based on the ASTM E122-17[18] standard and the selection of standards for AM was derived from the NISTIR 8005[19] technical report. Since a limited number of standards are available for AM, finding the equivalent conventional standard for additively manufactured test coupons was important. For the investigation of fatigue properties, the ASTM E0466-15 standard was selected, and design was also observed from the standard.Based on the statistical method described in the flow chart as well as budget limitations, sample counts were decided as shown in Table 2. For this study direction of the build was considered as another variable.

For this study direction of the build was considered as another variable. Therefore, X and Z direction samples along with four different post-processing conditions were considered for the experimental campaign to understand the influence on the fatigue properties.Experimental PlanMetal printing & parameters:Generally, maraging steels are low carbon steel with high strength and ductility. In this research EOS MS1 maraging steel powder was used. The chemical composition was derived by combustion-infrared absorbance and is mentioned in Table 1.

In this research EOS MS1 maraging steel powder was used. The chemical composition was derived by combustion-infrared absorbance and is mentioned in Table 1. Also, the particle size of the powder was in the range of 20 microns to 47 microns.Table 1 Chemical composition of EOS MS1(Reproduced with permission [17])Cu CaSaNi 18.37 9.07 4.97 0.67 0.09 0.18 0.02 0.007 0.04 0.05 <0.01 0.004 0.06 baseCo Mo TiAlCrMn SiPVChemical compositions mentioned in weight percentage; a -components were determined by combustion-infrared absorbance834FeAs mentioned, maraging steel fatigue samples were manufactured on an EOS M 280 machine. Initially, CAD files were prepared for the X and Z direction build. The preprocessing software for the machine produced the support structures. For this research, the block-type hatch support structure is used.

The preprocessing software for the machine produced the support structures. For this research, the block-type hatch support structure is used. For this research, the “performance” parameter set was selected from the EOS parameter sheet[20].Figure 2(a) top view of X-direction printed fatigue samples after the heat treatment (b) top view of Z-direction printed fatigue samples along with other test samples (without heat treatment)Samples were printed in X, Z direction and shown in Fig.2. Samples were printed for as- printed(A) condition and removed from the build plate by using wire cutting by Electrical Discharge Machining (EDM). For condition B, specimens were designed to be larger than normal so that 1mm of material could be removed from specimen surfaces by turning. These samples were printed, removed from the build plate by wire cutting, and further machined. For condition C, samples were printed from the machine and placed into the furnace before removal from the build plate.

For condition C, samples were printed from the machine and placed into the furnace before removal from the build plate. Similarly, for condition D, samples were printed, heat-treated, removed from the build plate by wire cutting, and further machined. These four different post-processing conditions were considered for this study and elaborated in Table 2 as an experimental plan.Table 2 Experimental planDirection of buildPrinted(A)Printed & Machined (B)Heat treated(C) Heat treated & machined (D)X Z20 2020 2010 1020 20Fatigue testingFatigue specimen geometry was derived from the ASTM E0466-15 standard and mentioned in Fig.3. During the design process, support structure height and wire cutting machine’s835wire thickness should be considered to avoid damage to the specimen during the wire cutting. For this research 1mm and more were added to support structure height.

For this research 1mm and more were added to support structure height. Similarly, specimens selected for machined conditions (B, D) had a 1 mm machining (turning) allowance added during design. The selected test method was for axial fatigue loading and especially the high cycle fatigue (HCF) method was followed.Figure 3 Specimen geometry for fatigue testFatigue tests were performed in IMR Test Labs, Singapore. MTS 810(model 318.1), MTS 370(model370.1), Instron 3(model no 8801) frames with different load capacities were used to perform the fatigue testing. For this study, a fully reversed loading cycle was applied. Based on the fully reversed axial loading S-N curve was formed with stress amplitude and number cycles (log scale) in the Y and X-axis.Results & discussionsFor the fatigue test ideally, a large number of test samples is required. An experimental campaign was designed with 10-20 and more samples for each condition.

An experimental campaign was designed with 10-20 and more samples for each condition. Fatigue tests were performed according to the ASTM standard with reversed cycle conditions. Usually, a fully reversed 𝜎𝑚𝑎𝑥(𝑚𝑎𝑥𝑖𝑚𝑢𝑚 𝑠𝑡𝑟𝑒𝑠𝑠 𝑜𝑟 𝑡𝑒𝑛𝑠𝑖𝑙𝑒) = −𝜎𝑚𝑖𝑛(𝑚𝑖𝑛𝑖𝑚𝑢𝑚 𝑠𝑡𝑟𝑒𝑠𝑠 𝑜𝑟 𝑐𝑜𝑚𝑝𝑟𝑒𝑠𝑠𝑖𝑣𝑒), stress ratio: R=-1, Amplitude ratio: A= infinity. Samples were placed in grips and push-pull axial loads were applied for the fatigue test at room temperature. HCF regime of the test was considered for the analysis.cyclefollows836PrintedPrinted & Machining Heat treated Heat treated & Machining120011001000) a p M900800( e d u t i lp m A s s e r t S70060050040030020010001.00E+031.00E+041.00E+05 Life cycles1.00E+061.00E+07Figure 4 S-N curves for the samples printed in the X-direction.Since it was an HCF test, the “Stress-lifecycle (S-N) approach” fatigue failure model was used for this study.

But using this method, distinguishing crack initiation and propagations was difficult. To define further about S-N approach, the equations listed below were utilized in this research.𝑐𝑦𝑐𝑙𝑖𝑐 𝑠𝑡𝑟𝑒𝑠𝑠 𝑟𝑎𝑛𝑔𝑒: ∆𝜎 = 𝜎𝑚𝑎𝑥 − 𝜎𝑚𝑖𝑛 − − −Equation 1𝑆𝑡𝑟𝑒𝑠𝑠 𝑎𝑚𝑝𝑙𝑖𝑡𝑢𝑑𝑒: 𝜎𝑎 =𝜎𝑚𝑎𝑥−𝜎𝑚𝑖𝑛 2− − −Equation 2𝑀𝑒𝑎𝑛 𝑠𝑡𝑟𝑒𝑠𝑠: 𝜎𝑚 =(𝜎𝑚𝑎𝑥+𝜎𝑚𝑖𝑛) 2− − −Equation 3𝑆𝑡𝑟𝑒𝑠𝑠 𝑟𝑎𝑡𝑖𝑜: 𝑅 =𝜎𝑚𝑖𝑛 𝜎𝑚𝑎𝑥− − −Equation 4𝐴𝑚𝑝𝑙𝑖𝑡𝑢𝑑𝑒: 𝐴 =𝜎𝑎 𝜎𝑚− − −Equation 5𝜎𝑚𝑎𝑥 is different for each post-processing condition’s fatigue test. Also, the cyclic stress range for the fatigue test was reduced to understand the life cycle of the material. Once the fatigue test was performed in the MTS machines, load data points were converted as stress values and plotted in the S-N curve. By using equation 1 stress range can be predicted and equation 2 gave values for stress amplitude. Similarly, Equations 3-5 can be used to find the mean stress, stress837ratio, and amplitude, respectively.

Similarly, Equations 3-5 can be used to find the mean stress, stress837ratio, and amplitude, respectively. Stress amplitudes and life cycles were plotted (in log scale on the X-axis). In fig 4, all four post-processing conditions of X-direction printed samples are plotted. For the various condition, the same load (490 MPa) different S-N curves were observed, however considerable scatter was evident. Importantly, it was observed that machining improved the fatigue strength significantly. S-N curves were shifted above the printed samples’ fatigue value. The rough surface of printed samples likely lowered fatigue life compared to the smoother after machining. Around 10 million cycles' fatigue value is significantly improved for condition B.

The rough surface of printed samples likely lowered fatigue life compared to the smoother after machining. Around 10 million cycles' fatigue value is significantly improved for condition B. Also, machining followed by micro shot-peening was recommended in the literature [21] for improved fatigue strength because it improves the compressive residual stress and closes the pores over the surface.PrintedPrinted & Machining Heat treated Heat treated & Machining120011001000) a p M900800( e d u t i lp m A s s e r t S70060050040030020010001.00E+031.00E+041.00E+05 Life Cycles1.00E+061.00E+07Figure 5 S-N curves for the samples printed in the Z direction.Similarly, heat-treated X-direction printed samples had improved fatigue strength but were not high as machining. During aging, material strength increased (approximately doubles compared to as printed) and residual stresses were relieved.

During aging, material strength increased (approximately doubles compared to as printed) and residual stresses were relieved. However, aging reduces ductility which likely reduces fatigue life. When machining and heat treatment processes were combined, fatigue strengths improved significantly because heat treatment increased strength and machining improve the surface finish.Z-direction printed samples exhibited lower fatigue strength compared to X-direction printed samples for all post-processing conditions. The main reason for the lower fatigue life was the anisotropy inherent in the layer-wise PBF process. So, based on the results, build directions838have a significant role to play in the fatigue strength. Other than that, Z-direction printed samples followed the same trends and can be observed in fig 5. Similar to X-direction S-N curves, machining played a significant role in improving fatigue life.

Other than that, Z-direction printed samples followed the same trends and can be observed in fig 5. Similar to X-direction S-N curves, machining played a significant role in improving fatigue life. Heat-treatment alone had a modest effect, but heat-treatment and further machining improved the fatigue life significantly.Despite the scatter of AM results in the S-N curve, it was observed that build orientation has a significant role in the fatigue behavior of the maraging steels. Similarly, each post-processing condition influenced the fatigue strength of the material. Especially, machining and heat treatment combined with machining played a significant role in both X and Z directions.ConclusionThis research study investigated the fatigue strength of maraging steel manufactured by an EOS M280 PBF machine for different post-processing conditions such as as-printed(A), printed and machined(B), heat-treated(C), and heat-treated and machined(D) along with X and Z built orientations.

A characterization method for additive manufactured materials was used to select the sample count and select equivalent test methods for fatigue tests. S-N curves were developed for tests conducted using the HCF regime. The findings of the research study are summarized below. 1. X-direction printed samples without any post-processing provided better fatigue strength than Z-direction printed samples. The main reason was the anisotropy inherent in the PBF process. Therefore, the build orientation influenced the fatigue life of the samples.2. Surface defects and high surface roughness were reduced by machining, which improved fatigue strength and life significantly.3. Aging heat treatment improved the fatigue life but not as significantly as machining. 4. Aging and machining resulted in significantly improved fatigue strength, compared to the as-printed condition. This combination resulted in the highest fatigue strengths and lives and should be recommended for in applications.

This combination resulted in the highest fatigue strengths and lives and should be recommended for in applications. AcknowledgmentThe authors would like to thank the funding agency, NAMIC (National Additive Manufacturing and Innovation Center), funded by the Singapore National Research Foundation, through grants 2017295 and 2018240, as well as partners ST Engineering Land Systems and the Singapore Armed Forces, for their in-kind contributions. The authors acknowledge the efforts of Senior Specialist Mr. Tan Kai Lee, Digital Manufacturing and Design Centre, SUTD, Singapore, for assistance with specimen fabrication and testing.References[1][2][3]Ian Gibson; David Rosen; Brent Stucker; Mahyar Khorasani, Additive Manufacturing Technologies, 3rd edition. Springer, 2021. J. Suryawanshi, K. G. Prashanth, and U. Ramamurty, “Tensile, fracture, and fatigue crack growth properties of a 3D printed maraging steel through selective laser melting,” J. Alloys Compd., vol. 725, pp.

tigue strength prediction through defects-based analysis of L-PBF 17-4 PH stainless steelIndrajit Nandi1,2, Jade Welsh3, Jutima Simsiriwong*1,3, Shuai Shao1,2, Nima Shamsaei1,21 National Center for Additive Manufacturing Excellence (NCAME), Auburn University, Auburn, AL 36849, USA 2 Department of Mechanical Engineering, Auburn University, Auburn, AL 36849, USA 3 School of Engineering, University of North Florida, Jacksonville, FL 32224, USACorresponding author: Email: j.simsiriwong@unf.edu Phone: (904) 620-5351AbstractMurakami’s approach has been used in the high cycle fatigue regime to relate the fatigue limit to the critical defect size and location in additively manufactured (AM) metallic materials. However, the applicability of this model has not yet been thoroughly examined for AM materials in the very high cycle fatigue (VHCF) regime.

However, the applicability of this model has not yet been thoroughly examined for AM materials in the very high cycle fatigue (VHCF) regime. Therefore, this study investigates the possibility of relating the volumetric defect features to the fatigue strength of 17-4 precipitation hardened (PH) stainless steel (SS) manufactured via laser-powder bed fusion (L-PBF) additive manufacturing technology. The 17-4 PH SS specimens are manufactured using an EOS M290 L-PBF system, heat-treated, machined, polished, and tested in the VHCF regime using an ultrasonic fatigue testing system.

The 17-4 PH SS specimens are manufactured using an EOS M290 L-PBF system, heat-treated, machined, polished, and tested in the VHCF regime using an ultrasonic fatigue testing system. Careful fractography has also been performed on all fractured specimens to determine the volumetric defects responsible for the crack initiation.Keywords: Murakami’s model; Laser powder bed fusion; Very high cycle fatigue; 17-4precipitation hardened stainless steelIntroductionAdditive manufacturing (AM) processes for metallic materials, such as laser-powder bed fusion (L-PBF), are becoming more widely used for part fabrication across various industry sectors [1]. This is due to many advantages, particularly its ability to fabricate parts with intricate geometries and net-shaped conditions that cannot be achieved with conventional manufacturing techniques.

This is due to many advantages, particularly its ability to fabricate parts with intricate geometries and net-shaped conditions that cannot be achieved with conventional manufacturing techniques. For AM parts to be fully adopted in safety-critical applications, it is essential to understand the mechanical behavior of AM parts, especially under cyclic loading, to obtain the basis of design allowable for a given application. The fatigue behavior of AM parts is known to be significantly influenced by the formed microstructure and AM process-induced defects [2,3]. Defect’s shape, size, and locations affect the stress intensity factor (SIF) surrounding the defects and microstructure, which influences the crack initiation behavior during fatigue loading [4].In some applications, stainless steel (SS) AM parts, such as those in gas turbines and engines, are subjected to cyclic loading at a very high frequency.

The expected service life of these parts can surpass ten million cycles (~Nf > 107) into the very high cycle fatigue (VHCF) regime. In the VHCF regime, where the crack initiation stage is significantly dominant relative to the crack438growth stage, the fatigue resistance of AM parts with the machined surface finish is primarily controlled by the process-induced internal defects [5]. By utilizing the defects characteristics (i.e., defects location and size), a fracture mechanics-based defect-sensitive √𝑎𝑟𝑒𝑎 model, proposed by Murakami [6] has been widely employed to predict the fatigue limit of AM parts, especially in the high cycle fatigue (HCF) regime [7–9].

However, the applicability of this model has not yet been fully examined for AM materials in the VHCF regime.17-4 precipitation hardened (PH) stainless steel (SS) is a widely used martensitic/austenitic grade SS due to its excellent combination of corrosion resistance, high strength, and tensile toughness, depending on the post-processing heat treatments, making it a suitable material to adopt for the AM processes to fabricate intricate parts [10]. However, the fatigue behavior of 17-4 PH SS is known to be more sensitive to defects, such as pores or inclusions, compared to other metals available for AM [11]. For wrought 17-4 PH SS, fatigue crack initiation in the VHCF regime is typically observed at nonmetallic inclusions or defects [8,12,13]. Nie and Mutoh [13] used the model developed by El Haddad et al. [14] to predict the fatigue limit of 17-4 PH SS by assuming the failure initiates from the material matrix. Schönbauer et al.

Nie and Mutoh [13] used the model developed by El Haddad et al. [14] to predict the fatigue limit of 17-4 PH SS by assuming the failure initiates from the material matrix. Schönbauer et al. [15] utilized Murakami’s approach to determining the fatigue limit for wrought 17-4 PH SS with varying stress ratios and defect sizes in the VHCF regime, and conservative predictions were obtained. They attributed subsurface inclusion (𝐴𝑙2𝑂3) as the fatigue crack initiator at lower stress levels, whereas crack initiated at the surface at higher stress levels.In this study, the applicability of Murakami’s model for fatigue strength prediction, for cycles to failure (~Nf > 107), in the VHCF regime for L-PBF 17-4 PH SS has been investigated. Ultrasonic fatigue tests were performed on heat-treated L-PBF 17-4 PH SS specimens fabricated using N2 as a shielding gas. Fracture surfaces were carefully analyzed to obtain the crack initiation mechanisms in all specimens.

Fracture surfaces were carefully analyzed to obtain the crack initiation mechanisms in all specimens. The defect characteristics, including the type, location, and size, were analyzed and utilized in the Murakami’s model. Finally, the fatigue strength prediction of L- PBF 17-4 PH SS specimens in the VHCF regime were presented and discussed.Materials and MethodsSpecimen FabricationCylindrical 17-4 PH SS bars specimens were fabricated using EOS M290, an L-PBF system, with Ar-atomized 17-4 PH SS powder and N2 as a shielding gas. The EOS recommended process parameters for 17-4 PH SS were utilized in this study and listed in Table 1. These cylindrical specimens with a height of 75mm and a diameter of 12 mm were fabricated in the vertical direction (z-direction), as shown in Figure 1(a).Table 1.

These cylindrical specimens with a height of 75mm and a diameter of 12 mm were fabricated in the vertical direction (z-direction), as shown in Figure 1(a).Table 1. L-PBF process parameters for 17-4 PH SS used in this studyPower (W)Scan Speed (mm/s) Hatching Distance (µm) Layer Thickness (µm)22075510040439Following the fabrication, all specimens were heat-treated according to ASTM A693 standard [16] to CA-H1025 condition, which consists of Condition A (CA) (1050°C/0.5 hour/air- cooled) with one step aging H1025 (552 °C/4 hours/air-cooled), as illustrated in Figure 1(b). After the heat treatment, the cylindrical specimens were machined into final geometries for ultrasonic fatigue testing, as shown in Figure 1(c). All L-PBF 17-4 PH SS specimens were manually polished using 280 grit, 500 grit, 1000 grit, 2000 grit, and 4000 grit silicon carbide papers and cleansed to remove the machining marks from gage section prior to the mechanical testing.Figure 1.

(a) Build direction of the cylindrical specimen, (b) applied heat treatment processes, and (c) final machined geometry of the ultrasonic specimen. All the dimensions are in mm.Fatigue testingUltrasonic fatigue tests were performed using a Shimadzu USF-2000A ultrasonic fatigue testing machine with uniaxial fully-reversed (R= -1) loading condition at a 20 kHz test frequency [17]. Internal heat generation in the specimen due to very high frequency loading was minimized by using an intermittent pulse/pause time of 110 msec and 800 msec oscillations and compressed cooled air. Specimen failures were defined when the resonance frequency of the specimen changed by ±500Hz. Run-out specimens were specified for specimens with no failure after 109 cycles. All tests were performed at room temperature and relative humidity.

Run-out specimens were specified for specimens with no failure after 109 cycles. All tests were performed at room temperature and relative humidity. Following the ultrasonic fatigue test, fractography analysis was performed for all L-PBF 17-4 PH SS specimens using a Tescan VEGA3 scanning electron microscope (SEM) to investigate the fracture mechanism and crack initiation sites. Before imaging, all the fracture surfaces were cleansed with methanol to remove dirt in a sonication bath.Defect-based modeling for fatigue strength predictionThe local stress field around the defects plays a crucial role in fatigue crack initiation and propagation. According to Murakami and Endo [18], the fatigue limit of a material can be440determined based on the threshold stress for a small crack around the defect.

According to Murakami and Endo [18], the fatigue limit of a material can be440determined based on the threshold stress for a small crack around the defect. The maximum SIF, 𝐾𝑚𝑎𝑥, of the crack around the defects for an applied uniaxial tensile stress, 𝜎, can be expressed as𝐾𝑚𝑎𝑥 = 𝐶𝜎√𝜋√𝑎𝑟𝑒𝑎 where the parameter C is 0.5 and 0.65 for surface defects and subsurface/internal defects, respectively. Since defects are usually irregularly shaped, a representative dimension of the crack or √𝑎𝑟𝑒𝑎 is used in this model, which is the area projection of the defects onto the plane perpendicular to the applied stress direction, as illustrated in Table 2.

Moreover, the threshold condition for non-propagating cracks can be expressed as follows [19]:∆𝐾𝑡ℎ = 𝐶1(𝐶2 + 𝐻𝑉)(√𝑎𝑟𝑒𝑎)1/3where 𝐶1and 𝐶2 are material constants, and 𝐻𝑉 is the Vickers hardness in kgf/mm2.At the fatigue limit, when cracks propagating from a defect cease, the SIF range at the crack tip can be expressed as the threshold SIF range (∆𝐾 = ∆𝐾𝑡ℎ). As a result, the fatigue limit, 𝜎𝑤, of a material can be obtained using the following equation [20].𝜎𝑤 =𝑚(𝐻𝑉 + 120) 1/6(√𝑎𝑟𝑒𝑎) where the constant 𝑚 are 1.41, 1.43, and 1.56 for subsurface defect, surface defect, and internal defect, respectively, as displayed in Table 2. Following Nezhadfar et al. [21] study, subsurface defects are defined by the defects with a linear distance less than 40 µm from the closest surface, while internal defects are those located greater than 40 µm from the surface.It should be noted that the number of cycles to failure representing the fatigue limit is not specified in Eq. (3).

The results indicate that the fatigue life of L-PBF 17-4 PH SS ultrasonic specimens spans from low cycle to beyond 107 reversals to failure, i.e., into the VHCF regime, and the horizontal asymptote of the S-N curve to determine the classical fatigue limit is not observed. It should be noted that the detailed microstructure analysis as well as the statistical distribution of defects obtained via x-ray CT scan of these specimens are not in the scope of the current work and can be found in a different study [22]. Also, in Figure 2, the origin of fatigue cracks (i.e., surface/subsurface defect or internal defect) was also identified. The fatigue failure was observed to originate from defects in different regions, similar to other VHCF studies for wrought high-strength steels [23].

The fatigue failure was observed to originate from defects in different regions, similar to other VHCF studies for wrought high-strength steels [23]. As seen in Figure 2, as stress amplitude decreases, fatigue life increases into the VHCF regime where crack initiation transitions from surface/subsurface to the internal region of the specimen, which approximately occurs at 108 reversals to failure.22Figure 2. Ultrasonic fatigue-life data for L-PBF 17-4 PH SS specimens.442At lower stress levels, where fatigue failure originated from an internal defect, a fish eye area, which is a flat ring-like smooth region, are observed, as indicated in Figure 3. Inside the fish eye, there is a bright fine-grained area (fine granular topographic structure) around the crack initiating defect, which is also referred to as an optically dark area (ODA) or a fine granular area (FGA) [20,24,25].

The formation of ODA occurs around the internal defects in the VHCF regime, which has been attributed to the lower SIF of the cracks around the defect [15,20]. If the SIF of the cracks around the defects is higher than the threshold value, then ODA does not form neither the fish-eye [26]. ODA formation ends when the SIF reaches the threshold value for crack propagation inside the fish-eye area. For high-strength steels, Murakami et al. [27] attributed trapped hydrogen around the defects as the driver of microcrack nucleating from the defects [18] at lower stress levels, which continues up to the circumference of ODA until the crack becomes self-propagable. Shiozawa et al. [28] reported the decohesion of small carbides in the vicinity of defects/inclusions leading to the formation of small cracks, which coalesce with each other inside the ODA.Figure 3.

[28] reported the decohesion of small carbides in the vicinity of defects/inclusions leading to the formation of small cracks, which coalesce with each other inside the ODA.Figure 3. (a) Fatigue fracture surfaces of L-PBF 17-4 PH SS specimens tested at 475 MPa with 308,700,000 reversals to failure, (b) magnification of the fisheye, and (c) magnification of the ODA.The location of the crack initiating defects was plotted against the stress amplitude, as displayed in Figure 4. It can be observed that as the stress level decreases, the crack initiates further away from the surface of the specimen towards the center. The transition from surface/subsurface crack initiation (less than 10 µm from the surface) to internal crack initiation (up to 1,300 µm from the surface) occurs at approximately 500 MPa.443223Figure 4.

Stress amplitude against the distance of the crack initiating defects from the surface for L-PBF 17-4 PH SS specimens.Figure 5 displays the representative size of the crack initiating pores, √𝑎𝑟𝑒𝑎, against reversals to failure for all L-PBF 17-4 PH SS specimens. It can be observed in Figure 5a that for both internal and surface/subsurface defects, the fatigue life increases with the reduction in defect size. The size of ODA, formed around the internal defects, increases with the decrease in defect size, which can be attributed to the lower SIF of the crack inside the ODA. The larger area of ODA could be attributed to the small internal defect size and the corresponding threshold SIF required for the crack to be self-propagable within the fish eye region of the fracture surface [20,26].

For the VHCF regime, Murakami [20] reported that the use of √𝑎𝑟𝑒𝑎 of defect may lead to non- conservative prediction for several high-strength steels, and suggested that √𝑎𝑟𝑒𝑎 of ODA can be used in place of the √𝑎𝑟𝑒𝑎 of defect in Eq. (3). Therefore, the √𝑎𝑟𝑒𝑎 of ODA for L-PBF 17-4 PH SS specimens with internal failure were also obtained in this study, as shown in Figure 5b.444Figure 5. A representative size, √𝑎𝑟𝑒𝑎, of defect or ODA against reversals to failure for L-PBF 17-4 PH SS specimens.Using Eq. (4), the fatigue strength for L-PBF 17-4 PH SS specimens was obtained from either √𝑎𝑟𝑒𝑎 of crack initiating defect, or √𝑎𝑟𝑒𝑎 of the ODA surrounding the crack initiating defect. The analysis was performed for specimens whose lives are in the VHCF regime (i.e., Nf > 107). The comparison of the experimental and predictive fatigue strength with scatter bands of 5 and 10 is displayed in Figure 6. The experimental fatigue strength is the stress amplitude of individual fatigue testing.

The experimental fatigue strength is the stress amplitude of individual fatigue testing. As can be seen, using √𝑎𝑟𝑒𝑎 of defect provides a better correlation between the experimental and predicted fatigue strength, with 67% of the data are within scatter bands of 10. Since the expression for parameter 𝛽 in Eq. (4) was empirically determined from the data in Wang et al. [21] to include the dependency of fatigue life, the parameter 𝛽 espression may need to be optimized based on the fatigue data in the current study for 17-4 PH SS. In addition, in Eq. (4), the Vickers hardness of 300 kgf/mm2 was utilized for all analyses in this study. This hardness value was obtained at the surface of the L-PBF 17-4 PH SS specimen; however, the hardness may vary across the cross-section of the specimen, which would require further investigation [20].445Figure 6.

Predicted fatigue strength versus experimental fatigue strength in the VHCF regime with factors of 5 and 10 scatter bands for L-PBF 17-4 PH SS.ConclusionsIn this study, the Murakami’s √𝑎𝑟𝑒𝑎 approach was utilized to obtain the fatigue strength in the VHCF regime of L-PBF 17-4 PH SS. Ultrasonic fatigue tests were performed, and fractography analysis was conducted on all specimens to identify the crack initiating defects and failure region. The following conclusions can be drawn from this study:1. The fatigue life of L-PBF 17-4 PH SS ultrasonic specimens spans from low cycle to beyond 107 reversals to failure, and the horizontal asymptote of the S-N curve to determine the classical fatigue limit was not observed.2. All fatigue cracks initiated from gas entrapped pores in all the specimens. As the stress level lowered, the crack initiation sites shifted from the surface/subsurface region towards the internal region of the fracture surface.

As the stress level lowered, the crack initiation sites shifted from the surface/subsurface region towards the internal region of the fracture surface. This transition occurred approximately at 108 reversals to failure.3. For L-PBF 17-4 PH SS specimens with internal failure, the optically dark area (ODA) was observed surrounding the crack initiating pore. As the internal defect size becomes smaller, the size of ODA increases.4. The modified Murakami’s √𝑎𝑟𝑒𝑎 approach was adopted in this study to include the fatigue life dependency in the VHCF regime. The predictive fatigue strength using √𝑎𝑟𝑒𝑎 of defects correlated well to the experimental data for L-PBF 17-4 PH SS in this study.446AcknowledgmentThis paper is based upon the work partially funded by the National Institute of Standards and Technology (NIST) under Award No. 70NANB19H170. This paper describes objective technical results and analysis.

M of ABS Patterns for Investment CastingByPaul Blake* Eric Fodran"" Martin Koch"" Unny Menon"" Bill Priedeman*Stephen Sharp""STRATASYS"" CALIFORNIA POLYTECHNIC STATE UNIVERSITY, SAN LUIS OBISPOABSTRACT This paper will evaluate the suitability of fused deposition modeling (FDM) ofacrylonitrile butadiene styrene (ABS) patterns for use in investment casting. The focus is on integration with existing foundry practices. It is a combined industry/university project with the case studies performed at the industrial sites with university produced patterns. Process parameters, ash handling and casting issues are addressed.1952INTRODUCTION Fused deposition modeling (FDM) has demonstrated the ability to reduceproduction time requirements in many applications. Like other forms of rapid prototyping, FDM provides a means of directly creating a tangible 3D object from a CAD/CAM image.

Like other forms of rapid prototyping, FDM provides a means of directly creating a tangible 3D object from a CAD/CAM image. On average when compared to machining, FDM has provided a method of verifying designs quickly and often, for a low relative cost. Additional decreases in production time can be obtained, if the FDM Acrylinitrile butadiene styrene (ABS) builds were used as the investment casting patterns. The feasibility of this is determined by addressing casting issues, ash handling, and process parameters, as well as, through thermal analysis and macroscopic examination.3GOALS Our goals are to present casting guidelines for the average Investment Casting Foundry (IC) to use with the ABS patterns. These guidelines involve the use of simple techniques and common materials, and have generally resulted in successful castings.CASTING ISSUES In order for ABS casting patterns to reduce the investment casting time cycle, ABS must be easily useable with standard practices.

Therefore, pattern preparation, gating and venting, investing, burn-out, ash removal, and casting must be considered when determining if ABS is a usable alternative to wax.44.1Pattern Preparation The ABS patterns are robust and easy to work with. They sand well and can bemachined. They can be modified with wax to improve surface finish, and combined with waxes to form composite patterns.4.2Gating and Venting For ABS patterns, traditional gating and venting practices require somemodifications. These changes can easily be adapted to IC, and must be followed in order to obtain adequate ash removal. It is during the gating and venting design process that the modifications are first introduced.Ash removal will require low air pressure and water rinsing. Therefore, it isnecessary to install additional vents on the pattern. This will improve the air and water feature penetration during the ash removal process.

Therefore, it isnecessary to install additional vents on the pattern. This will improve the air and water feature penetration during the ash removal process. A fair amount of access to the interior of the mold structure is also gained by cutting open the gate end. This too will increase the efficiency and ease of the ash removal. Consequently, it is also important to design the investment tree with the idea that the gate and vent ends will be removed prior to the burn-out process.Once the investment tree design has been established, standard IC technique isused to attach the gates and venting.4.3Investing A cleaning/etching agent is normally applied to the investment tree prior toshelling. This provides a means of removing any release agents. Once the196cleaning/etching agent has dried, standard investment casting procedures are followed in producing the shell.

This provides a means of removing any release agents. Once the196cleaning/etching agent has dried, standard investment casting procedures are followed in producing the shell. Generally an FDM investment shell is composed of, first, three normal layers, followed by three layers with steel wire reinforcement, with a final seal coat then applied. The number of reinforcing layers can be increased to strengthen the shell when necessary.After the final seal coat has dried, the tips of the vents are ground to expose thewax. The main sprue is also ground exposing the wax. This modified investing step, must be performed prior to the burn-out cycle.4.4Burn-Out Two methods have demonstrated potential for removing Rapid Prototype (RP)patterns from an investment mold. Though both techniques have succeeded in reducing the pattern and associated material to ash, each has inherent drawbacks.The autoclave / high temperature burn method is a two step process.

Though both techniques have succeeded in reducing the pattern and associated material to ash, each has inherent drawbacks.The autoclave / high temperature burn method is a two step process. First, the mold is placed in an autoclave to remove the majority of the wax from the mold. This step is considered the melt-out cycle, with temperatures ranging between 200 OP to 300 OF (94°C to 150°C) (Jacobs 1996). The mold is then transferred, in the second step, to a high temperature oven. At this stage the pattern material is reduced to ash and any residual wax is incinerated. During this stage, the burn-out temperature will range between 1,200 of to 14000P (650°C to 760°C) (Jacobs 1996). A disadvantage with this method is that on occasion, the shelling has been reported to crack.

A disadvantage with this method is that on occasion, the shelling has been reported to crack. This is due to the expansion of the RP material during the melt-out cycle and is addressed by using additional layers and reinforcement..The temperature for the flash process, starts at 16000P (871°C) and is elevated to as high as 19000P (1038°C) for burn-out. This method reduces the RP pattern, as well as all associated tree material, to ash for easy removal. This method too has a drawback. All of the wax in this technique is incinerated, which creates additional smoke during the process.Both methods have worked successfully with FDM ABS molds.Foundries should determine which method is preferred for their casting process.4.5Ash Removal Removing the ash requires an additional step when casting with a FDM ABSpattern.

After the shell has been blown-out with low pressure air and flushed with water, the open gate end and vent openings are resealed with shell repair cement.4.6Casting Casting procedures for FDM ABS mold and traditional investment casting are thesame. Standard filling methods and furnace temperatures are applicable.CAST PATTERNS ABS is the pattern material to be studied. It is the name given to a polYmeric structure that represents a family of thermoplastic materials (Smith 1990). ABS is a mixture of three monomers: polyacrylonitrile, polybutadiene, and polystyrene. ABS is not a terpolYffier though, it is based on blended copolymers of stYrene-acrylonitrile5197(70:30) and butadiene-acrylonitrile (65:35) and on graft interpolymers of styrene and acrylonitrile with polybutadiene (Smith 1990) (Sharp 1990). By adding modifiers to the basic ABS mixture, the material properties can be altered.

By adding modifiers to the basic ABS mixture, the material properties can be altered. Consequently, the ABS material being evaluated should be the same material that is used during the production of a FDM build.THERMALANALYSffi Thermogravimeteric Analysis Systems (TGA) have proven effective for temperature related polymeric degradation studies. The TGA is designed to measure the change in the sample's mass with respect to the temperature of its environment, even if that environment is altered (Kelen 1983). Several methods can be used to alter the environment: varying the temperature, introducing reactive gases into the sample chamber or a combination of these variables.6After determining the appropriate test parameters, two samples are selected forexposure. The first specimen is the actual sample material; contrarily, the other is inert to the test environment.

The first specimen is the actual sample material; contrarily, the other is inert to the test environment. The inert sample, known as the reference sample, provides a reference point during thermal analysis.6.1Graphical Interpretation Most TGA systems also provide the derivative of the thermogravimeteric curve(DTG) and label the curve, the Differential Thermal Analysis (DTA). The DTA provides information on endothermic or exothermic reactions in the sample material. This is depicted by plotting the change in temperature measurements for the sample against the unaffected reference sample. The endothermic and exothermic points are then useful in determining the temperature at which deterioration of the material properties occurs (Kelen 1983).MACROSCOPIC EXAMINATION Macroscopic examination is one of the most basic forms of visual examination available. Often used as a prelude to microscopic studies, macroscopic inspection can also be used alone.

Often used as a prelude to microscopic studies, macroscopic inspection can also be used alone. In this case, inspection of the cast parts will be visually inspected, unless it is deemed necessary to examine a part in greater detail.78EXPERIMENTATION Several builds were produced using the Stratysis 1600. These builds were thenused as investment casting patterns and for thermal analysis. The two different types of builds produced were the 3D Systems "Nine-Wall" patterns, seen in figure 1, and the Stratysis "Turbine" pattern, seen in figure 2.CASTING Each sample requires some preparation work prior to casting. Sanding was not required, but a utility knife was used to shave plastic stringers and remove some of the ABS support material. A small screwdriver was also utilized to remove some of the ABS support material from the patterns. The final preparation requirement for each pattern8.1198was to fill the small surface openings with paste wax.

The final preparation requirement for each pattern8.1198was to fill the small surface openings with paste wax. Approximately 20 minutes was required to prepare each pattern for investing.Figure 1 - FDM Nine-Wall ABS patternFigure 2 - The as cast turbine pattern on the left and ABS turbine pattern on the rightSome of the patterns were provided to Aurora Engineering for use as castingpatterns. A design that would be capable of providing adequate access for ash removal was determined first. The investment tree design and venting arrangement are show in figure 3. The investment tree was composed of a styrofoam core with a wax coating.

The investment tree design and venting arrangement are show in figure 3. The investment tree was composed of a styrofoam core with a wax coating. Addition venting is also evident in figure 3:> from the increase in the bleeder wax:> that is positioned around the ABS pattern.Figure 3 - Nine-Wall pattern ready for investingFigure 4 - Invested pattern with gate end and vents openedThe entire investment tree was sprayed with isopropyl alcohol and allowed to dry,in preparation for the investing process. Once the investment tree had dried:> it was dipped in a coating solution a total of seven times: three normal coats, three coats with steel wire added, and then one final coat. The final invested pattern is shown in figure 4. The ends of the vents and gates have been removed to aid in subsequent ash removal. Burn-out was performed using the flash process.

The final invested pattern is shown in figure 4. The ends of the vents and gates have been removed to aid in subsequent ash removal. Burn-out was performed using the flash process. The molds were placed in a preheated199furnace at 1600OP(8710C), immediately raised to 19000P (1038°C) and left for 90 minutes. After burn-out, the ash (a light white powder) was remove with low-pressure air and water. Mold repair cement was then used to reseal the vent and gate ends.Standard casting procedures were then followed and aluminum was cast into themolds. The molds at the time of casting were preheated to 12000P (649OP) and the aluminum was poured at 12850P (696°C).The molds were allowed to solidify and the shelling was removed. This producedthe as cast structure that is seen in figure 5. After removing the excess material, the gating and venting contact points were ground smooth.

This producedthe as cast structure that is seen in figure 5. After removing the excess material, the gating and venting contact points were ground smooth. The finished products are visible in figure 6, as well as in the Turbine Pattern comparison photograph labeled figure 2.Figure 5 - As cast Nine-Wall prior to removal of excess materialFigure 6 - Final Nine-Wall castingThermal Analysis ABS Samples for thennal analysis were taken from a Nine-Wall pattern. The TGA test was then programmed to simulate the environment that the ABS would be exposed to, and react in, during the burn-out process. The ABS was subjected to a temperature range that started a 71.6 OP (22.0°C) and was raised to 1112'}' (600°C). The gas environment used during the entire test period was air. At the completion of the test, the computer generated graphs showing the weight as a function of temperature, and the derivative of that function.

At the completion of the test, the computer generated graphs showing the weight as a function of temperature, and the derivative of that function. This graph is illustrated in figure 7.8.2200Figure 7 - TGA results showing degradation and exothermic peaks9RESULTS AND DISCUSSION The FDM ABS builds all performed successfully as IC patterns. The ABSpatterns, which were built at an academic institution, were transferred to industry and easily assimilated into the IC process.Upon visual inspection of the final IC cast products, surface appearance wasfound to be acceptable, and an example of the surface is shown in figure 8. The surfaces in the cast products may be improved further, by increasing the time spent preparing the surfaces of the patterns. Fine, thin wall details, as in the fins on the Nine-Wall pattern and the Turbine pattern, as seen in figures 9 and 2 respectively, were accurately replicated. The only noticeable imperfection was found at the tip of one fin.

The only noticeable imperfection was found at the tip of one fin. This was most likely due to the fin cooling too quickly and the metal solidifying before it had completely filled this given area, a cold short.The TGA confirmed that the ABS material bum-out process is above the required temperature for ABS to degrade. As was evident from figure 7, degradation of the plastic occurs at approximately 797<>P (425°C). This is displayed by the exothermic reaction and is estimated from this figure.201Figure 8 - Magnified view of the same as cast surface. The left picture was photographed with a light macroscope and the picture on the right with a SEM at 22.8X.Figure 9 - As cast Nine-Wall on top and the ABS Nine-Wall pattern on the bottom.10CONCLUSION This industriaVuniversity joint project, addressed the IC requirements for FDM ofABS builds to be used as patterns in industry.

It was determined that modifications would be necessary, but that these changes are easily adaptable with present IC techniques. Therefore, FDM ABS patterns are suitable for use in investment casting facilities.11 Jacobs, P. F. (1996) Stereolithography and other RP&M Technologies, Society of Manufacturing Engineers, New YorkREFERENCESKelen, T. (1983) Polymer Degradation, Van Nostrand Reinhold Company Inc., New YorkSmith, W. F. (1990) Principals ofMaterials Science and Engineering, McGraw-Hill, Inc., New YorkSharp, D. W. (1990) Dictionary ojChemistry, Penguin Books, London202.

lM SysteDls and Local Adaptive SlicingJustin Tyberg and Jan HeIge B0hnMechanicalEngineering,•Virginia Tech Blacksburg, VA 24061-0238, U.S.A. teL: 540-231-3276, fax.: 540-231-9100, e-mail: bohn@vt.eduFabrication times for layered •• manufacturing. system~can. be significantly reduced by continuously maximizing the build layer thickness, wit~inthe bounds.. ofthesystelll'~ fabricationcap~bilities and the surface deviation tolerance derived from the<part surface .. geometry.•• This paper describes how parts and features within a build volume can be adaptively. sliced independently of each other such that each is· fabricated with distinct build layer thicknesses. This enables significant redllction in fabrication times relative to conventional adaptive build layer thickness. techniques.

This enables significant redllction in fabrication times relative to conventional adaptive build layer thickness. techniques. The· new approach has been implemented ona FDM 1600 rapid prototypingsystem, together with a revised system calibration to ensure smooth surface transitions between dissimilar build layer thicknesses.1IntroductionFused Deposition Modeling (FDM) is today· the second most common commercial layered manufacturing system. It fabricates parts by extruding molten wax or thermoplasticrnaterial through a smalLnozzle to form a thin bead or "road" that is deposited in a predetermined pattern to complete each build layer, bondingthe extrudate to adjacent and previously deposited roads .. The motion of thy extrusion system operates under three-dimensional computer •numerical· control (CNC). The extrusion. system consists .of a sl11all ralllextruder in which the spooled filament feedstock pushes molten material through the liquefier. with low. shear..

The extrusion. system consists .of a sl11all ralllextruder in which the spooled filament feedstock pushes molten material through the liquefier. with low. shear. . The feedstock is driven into the.extruderby counter·rotating rollers as needed under Cl\TC,•. \Vhere it.isheated. tORlllanually set temperature, TL • The extrudate is deposited within a build chamber holding a manually set temperature, Tc In the FDM 1600 rapid prototyping system, TL and Tc can be up to 300°C and 70°C, respectively.The most common build material used with FDM systems.is P400 ABS plastic (Stratasys, Inc.), and it is available il1 several stock colors, inclu~ingwhite, red, blue, green, yellow and black. This material is,in theFDM 1600 rapid prototyping system, typically extruded with TLand Tc setat270°C and 70°C, respectively, through a 0.012" (0.30mm) nozzle orifice to form roads that are 0.010"(0.25mm)thick and 0.020" (0.51 mm) wide, while being deposited at a rate of 0.8 in/s{2 cm/s)..

This corresponds .toa maximum standarddepositi?n rate oiO.6in3Jhr. (9.4 .cm3/hr), that is reduced by delays in startillgand. stoppi~g ofdeposition, m()tionfrom one build layer. to the nsxt, •and the fabrication of support strt1ctures ..~s needed. This particular combination of hardware, materials and processing parameters has beenprecisely calibrated by Stratas~s,Inc.to.enableFDM users to .. produce some of the most. dimensionally accurate •parts •in •.• the layered •manufacturing industry; and it has proven itself particularly •.useful in. the prototype production of functipnal and semi-functional to-be-injection-molded plastic parts.Aswithm?st layered 11lanufacturing systems, there are two areas in.whichJ3I)Msystems e (l)Iyd~cti?n ofbuildtillles,~n(Li(2) could <benefit. from further improvements.

from further improvements. l'hese ar improvement of part surface smoothnesses.These needs for improvement are drivenby>designand marketing personnel's desire to· access smooth functional prototype parts within minutes of a design change. This, however, is still beyond the capabilities of current layered manufacturing technologies: Today a prototype part takes hours or days to build-and hours of manual labor to finish-only. for it .to.• often· bedisP9sed •.•of within .minutes of. its .completion once the quick It is therefore desirable to (a) increase inspection or test it was intended for has been performed. the average material deposition rate to reduce. the build time, and (b) reduce the build layer thicknesses to produce smoother parts that require less finishing.These two objectives can be approached.

the build time, and (b) reduce the build layer thicknesses to produce smoother parts that require less finishing.These two objectives can be approached. by applying the basic principle of continuously varying the material depositionrate such that, at all times, the maximum material deposition rate is used that satisfies the.local surface deviationtolerance across the part being fabricated. This principle has been demonstrated byadaptivelyadjusting the build layer thicknesses such that they are maximized subject to meeting •surface deviation tolerances [1-5], and by increasing build layer thicknesses in the interior of the. part [4,6]. As these two approaches .we combined, significant improvements in material deposition rates can be realized; limited primarily by the capabilities of the hardware. For instance,.

part [4,6]. As these two approaches .we combined, significant improvements in material deposition rates can be realized; limited primarily by the capabilities of the hardware. For instance,. in the case of the FDM 1600 rapid prototyping systems building with P400ABS plastic and using the standard 0.012" (0.30 mm) nozzle orifice, the maximllm physical material deposition rate achievable isA.3in3/hr (70.8 cm3/hr). This corresponds to a 0.030" (0.76 mm) thick by 0.050" (1.27 mm) wide road being deposited at 0.8 iuls (2 cm/s), which is 7.5 timesthe standard rate. If, in addition, a 0.100" (2.54 mm) air spacing is inserted between the roadsinterior to the part, then the maximum virtual material deposition rate increases to 13 in3lhr (212 cm3/hr) , or 22.5 times the standard rate.

We therefore see the potential of fabricating parts by FDM in hours instead of days while achieving a surface smoothness equivalent to that of 0.003 0.005" (0.07 - 0.13 mm) uniform thickness build layers.This paper reviews recent efforts towards achieving these objectives, and, in particular, that of fabricating. smooth-surface parts by FDM· in the minimum amount of time. The following sections describe a significant improvement to previous adaptive slicing techniques [7,8], issues that surfaced when migrating this work from computer simulations to actual fabrication, and how these issues were resolved to produce the desired outcome; namely, fast, smooth-surface parts by FDM.2Local Adaptive SlicingConventional aqaptive slicing techniques [1-5] have demonstrated the potential benefit of adaptive slicing, but have not madeit out of the laboratory setting.

In particular, adaptive slicing has only been applied to simple geometries, one part at a time; a situation that differs substantially from most industry settings.where part geometries are..often complex and where multiple parts are often fabricated concurrently within the build envelope. Figure 1 illustrates· the consequence of It shows how applying conventional adaptive slicing techniques to a more .. complex scenario. decisions made to satisfy the surface deviation tolerance of one part or part-feature, in this .case the sphere, overrule the lesser J1eeds of the more simplistic block. As a result, the block is fabricated with layers that are needlessly thin and valuable fabrication time is wasted.Figure 2 illustrates a more desirable situation. that results from fabricating the SaIIle parts independently of one another. A method to achieve this independence has been developed and is described in detail elsewhere [7,8].

that results from fabricating the SaIIle parts independently of one another. A method to achieve this independence has been developed and is described in detail elsewhere [7,8]. Briefly, this new approach, which has been termed local adaptive slicing, consists ofa three-step process that is based on stepwise uniform refinement [4,5]. First, all the parts in the build volume are sliced into thick uniform-thickness slabs. Then the contours on the top and the bottom of each slab at a given height are grouped into sub-slabs that376Part with complex surface i i I I Ik:~~.-~.~_Figure 1 The potential savings of adaptive slicing is lost by conventional slicing methods that slice all parts of a given build volume with the same resolution, regardless of their dissimilar surface characteristics.

In this figure, the thin layers necessary for the sphere are imposed unnecessarily on the block.Sphere is fabricated with thin I whereayers 1 neededBlock only requires thick layers<r""i '--IzL::1'.-... :..c.....,. _<!.f--'.- ----~ .._-l.£J--l jJxFigure 2 The new approach fabricates individual parts and part-features independently with distinct layer resolutions applied locally as necessary.reflect that these contours would be physically connected by a sllrface even if only the sub-slabs at that particular height were fabricated. Finally, each sub-slab is sliced independently into some integral number of uniform-thickness slices to satisfy. _the surface deviation •tolerance •of that particular sub-slab.

Finally, each sub-slab is sliced independently into some integral number of uniform-thickness slices to satisfy. _the surface deviation •tolerance •of that particular sub-slab. Hence, all sub-slabs are of equal thickness, regardless of how each one is subdivided into thinner build layers, and can therefore easily be processed and fabricated independently of one another.In the FDM 1600 rapid prototyping system, the build .volume .issubdivided into O.030~_' (0.76 mm) thick slabs, that correspond to the thickness ofthe maximum flow rate roads described in Section 1. Tlrisfacilitates the fabrication of the interior of each sub-slab using these roads and the exterior using thinner roads_ that _satisfy local surface -deviation -. tolerance requirements.[4,6]. Each sub-slabs' exterior regions are _subdivided into an integral.number of uniform thickness build layers, as/abO using the criteria developed by Dolenc and MakeHi [1] and.refined by Sabourinet at.

Each sub-slabs' exterior regions are _subdivided into an integral.number of uniform thickness build layers, as/abO using the criteria developed by Dolenc and MakeHi [1] and.refined by Sabourinet at. [4,5}, but limiting the points of consideration to only those thatare associated with a particular sub slab whelldetermining.•theoptimal subdivisionforthatsub-slaq:{}J asub-slab == mt--max nz( Lmax ~~.'asub-slab E[1] ,amax,. t( LmaxJa max - m L. mn377(1)where Lmin and Lmax areth~lniniIIlurp. and maximum build layel'thicknesses available, respectively; Cmax~Lmin is the maximumperrnissible surface deviation error; and {nz} is the set of unit normal z componentsfor all the points along all the contours of thesub..slab.

The software developed reads any CAD model describedinthe.STL fileforIIlat,~~neratecsthecontourson each build layer, and exports them as an .SSL file completewithlayertl1icknessandflow rate information. QuickSlice 5.0, the FDM post-processor by Stratasys, Inc., can then use this information generate the actual roads that are described in the .SML file formatfor control of the FDM 1600.The software appearsto be robust and it generates .SMLfileswith values that are identical to what would be generated if this process had been executedlTIanually. Unfortunately, the initial fabricated parts were not satisfactory. While the part surfaces where·the build layers were constant were smooth, including across suh..slabs, there was a small but noticeable discontinuity wherever there was a change in build layerthickness,andparticularly when there was a change between the 0.005" (0.13mrn)andO.015" (0.38 mrn)build layers.

In addition, the thinner layers tended to delaminate andthus further reducedthe resulting surface quality.~he smooth part surfaces where the •layer .thickness was constant··indicated that it is within the capability ofthe hardware and material to consistentlyproduce high-quality part surfaces. This led us to suspect. that the discontinuity problem was a function of poor calibration. .. A. closer examination of the calibration tables used by QuickSlice 5.0 strengthened this hypothesis. As shown in Figure 3, the ••• calibration curves. for several of the layer tl1icknessesused, are composed of several CO disconti~uous,piecewise logarithrniccurve segments. This is contrary to whatone would expect since there is no discontin~ity •• or sharp chan~e.i~\the flow characteristics of ABS plastic at th~seflow, pressure, and temperature levels... Therefore, if the .. transitions along a constantlayer-thickness curve are.

Therefore, if the .. transitions along a constantlayer-thickness curve are. impr~sise, ·the tra~~itions from .onebuild-layer-thickness curve· totheinext .would also haye tobeiml'recise,andthislatter •. i111precision ... would explain the discontinuities that were. experienced. on the part surfaces \Vh~n. c~angingJrom one buildlayer thickness to another. ThestandardFDM.calibration tables forP400 ABS plastic should therefore be revisedto accurately reflect actual material behavior so smooth, adaptively sliced part surfaces can be fabricated.The delarnination problem, on the otherhand, was not due to inaccurate calibration, but to insufficient· material weld time. The low material volumes and flow rates· associated with thin(2)200180U 160 CI> J!! 140 C :::: :IE 120 c..a:: 100-S 80~60DISCONTINUITIES: piecewise logarithmic curve segments\..............

This weakened weld often breaks under the pulling force exerted by the liquefier as it traces across the build plane. When this occurs, the newly deposited road will be free to lean in the direction of the motion of the liquefier, which .will change the road's location and disrupt the steady-state material flow rate, and, ultimately, distort the part geometry.The effective weld time can be extencled, to increase the strength ofthe weld, by increasin~ the extrusion temperature, TL• Stratasys, Inc. recommends increasingTL to 290°C andJowering Tc to 50°C whenfabricatingwithO.OOT' (0.18mm) build layers in. the FDM2000[91. We therefore duplicatedthese temperature changes in the.RDM1600, and we found that it effectively eliminated the delamination problem while still allowing fabrication with the thicker build layers.

We therefore duplicatedthese temperature changes in the.RDM1600, and we found that it effectively eliminated the delamination problem while still allowing fabrication with the thicker build layers. This observationis important because the temperature settings are set manually, and, therefore, all build layer thicknesses must be fabricated at the same temperature configuration. However, since the. material flow rate is affected by changes in temperature, •• and· the stanclard calibration tables.

However, since the. material flow rate is affected by changes in temperature, •• and· the stanclard calibration tables. are based on a different XL / Tc combination, achangein processing temperaturesca.lled for a revision of the ca.librationtables, at the new TL / Tc combination,to enable accurateandsmoothpart...surface fabrication.4CalibrationsBoth the apparent inaccuracies of the standard FDM 1600 ca.librationtablesandtl1echange inprocessillg temperatures made revising thtca.librationtables necessary to facilitate accurate andresultfrom variousc?nstantbuildlayer thickness .1 material feedrate/TL / .Tccombinationsihad to be measured.••·••These m.easurements were obtainedbybuildillgsevera.l. series. of vertical .•• walls. Each wall was. 2.0" {50 nun) long, 0.25"(6.4 ••• nun) tall, hadthe \width ofa single bead, and was assigned a flow rate value of even number between 2 and 254 (RPMlinch/sec)[10]. Eachseries379100160TL = 290°C Tc = 50~C'if 140 ~:i 120Q.

Eachseries379100160TL = 290°C Tc = 50~C'if 140 ~:i 120Q. ~ 100 ~ u:00R2 = 0.984R2 = 0.988.. • 0.0150 inch • 0.0075 inch0.0100 inch0.0050 inch60R2 = 0.981 ..40 -1 - - - -1 - - - - 1 - - - - - - 1 - - - -1 -__-1-__-1-__-1-__-1-__0.0120.0140.0160.Q180.0200.0220.0240.0260.0280.030Road Width (in)Figure 4 Revised FDM 1600 calibration v&lues for P400 ABS plastic extruded at TL =290·C through a 0.012" (0.30 mm) nozzle orifice into a Tc =50·C build envelope. Continuous logarithmic curves are fitted to data points that represent measured road widths for given flow rates at fixed build layer thicknesses.contained up to 21 parallel walls that were given sequentially increasing flow rates and were located about 0.4" (10 nun) apart on a single support material base.

This base was placed in the center of the build chamber, and its walls were oriented in the direction of the heated air flow in the build chamber, to provide near uniform material cooling.12 series of walls were fabricated using a nozzle orifice diameter of 0.012" (0.30 mm); TLTc settings of 270°C / 70°C, 290°C / 70°C, and 290°C.I 50°C; and build layer thicknesses of 0.0050", 0.0075", 0.0100", and 0.0150" (0.13, 0.19, 0.25, and 0.38 mm). The flow rates comprised all even numbers between 50 and 90, 70 and 110, 112 and 140, and 130 and 170, for the 0.0050", 0.0075" 0.0100", and 0.0150" (0.13, 0.19, 0.25, and 0.38 nun) build layer thickness, respectively. The head speed was kept constant at 0.8 in/s (20.3 mm/s). The external ambient conditions were measured with a thermometer and hydrometer on top of the FDM 1600.

The head speed was kept constant at 0.8 in/s (20.3 mm/s). The external ambient conditions were measured with a thermometer and hydrometer on top of the FDM 1600. The temperature measured between 78°F and 80°F (25°C and 27°C), and the relative humidity measured between 25% and 32%.Five measurements were taken from each wall using a caliper with a 0.0005" (0.01 nun) resolution. All measurements were obtained from the central regions of the lines to ensure that only steady-state deposition was being measured.. These measurements were graphed using Microsoft Excel 7.0 and fitted with a single continuous logarithmic curve for each build layer / Tc combination (Figure 4). The functions described by these curves where then thickness / TL used to revise the calibration tables used by QuickSlice 5.0.Results With the revised calibration tables installed, the part surface quality produced by the FDM 1600 rapid prototyping system noticeably improved.

The former discontinuities that occurred at transitions between dissimilar build layer thicknesses can no longer be observed visually or by human touch (ref. arrow in Figure 5). The only remaining sign of these transitions are the line density changes reflecting off the part surfaces.5380/Slower than if built C'Jf)n.nrnrJf)'"(a)(b)Figure 5 Two parts with a number of individual features were fabricated concurrently in a single build volume using (a) conventional adaptive slicing and (b) local adaptive slicing. Both sets hold the same overall surface tolerance, but set (b) requires 37% less fabrication time. The medium, dark, and light gray areas are 0.0050",0.0075", and 0.0150" (0.13, 0.19, and 0.38 mm) build layers, respectively. The arrow shows the location of the approximate 0.005" (0.13 mm) discontinuity before revising the calibration tables.The potential time savings from adaptively slicing each part-feature independently are substantial.

In Figure 5, the two parts were adaptively sliced with Cmax set to 0.0035" (0.09 mm) and fabricated with 0.0050", 0.0075", and 0.0150" (0.13, 0.19, and 0.38 mm) build layer thicknesses. With this configuration, local adaptive slicing reduced the fabrication time by 37% This. case illustrates how the performance of compared to conventional adaptive slicing. conventional adaptive slicing degrades as the number of parts and part-features increase: InFigure 5(a), the right-most "finger" of the left part requires thin build layers, and these thin layers are imposed onto the two other fingers and the vertical walls. on the part to the right. Hence, when using conventional adaptive slicing, it takes 12% longer to fabricate the two parts concurrently than sequentially. Local adaptive slicing, on the other hand, does not carry a penalty for.

Local adaptive slicing, on the other hand, does not carry a penalty for. concurrent fabrication; instead, in the case shown here, there is a 4% reduction because fewer verticallayer-to layer movements are needed during concurrent fabrication.6ConclusionsA new effective approach to adaptive slicing has been developed. It fabricates all parts and part-features independently of one another. Therefore, its effectiveness does. not deteriorate as part complexities and build volumes increase. Conventional adaptive slicing techniques, on the other hand, do deteriorate under such conditions, and are therefore impractical in an industriaLsetting.To implement local adaptive slicing onFDM 1600 rapid. prototyping system, it .was necessary to (1) increase the extrusion temperature to prevent delamination~hen fabricating with thin build layers, and (2) revise the calibration tables to enable accurate and smoothpart"surfaces.AcknowledgementFDM® is a registered trademark of Stratasys, Inc.

of Minneapolis, Minnesota, U.S.A.,Reg. No. 1,663,961.381References[1][2][3]P. Kulkarni and D. Dutta, "An.Acc~rateSlicing Computer-Aided Design, vol. 28, no. 9, September 1996, pp.Manufacturing,"[4]E. Sabourin, Interior, Layered Manufacturing,"M..S.thesis, Mechanical Engineering, Virginia Tech, Blacksburg, Virginia, U.S.A., February 1996. URL: http://cadserv.cadlab.vt.edu/bohn/projects/sabourin/"Adaptive High-Precision Exterior, •• High-Speed[5]E.Sabourin, S. A. Houser, and J. Refinement," Rapid Prototyping Journal, vol. 2, no. 4, 1996, pp. 20-26.B~hn, "Adaptive Slicing Using Stepwise Uniform[6]E.Sabourin, S. A. Houser,. and J. H. B~hn, "Accurate Exterior, Fast Interior Layered Manufacturing," Rapid Prototyping Journal, vol. 3, no. 2, 1997, pp. 44-52.[7]J. Tyberg, "Local Adaptive Slicing for Layered Manufacturing," M.S. thesis, Mechanical Engineering, Virginia Tech, Blacksburg, Virginia, U.S.A., February 1998.

3, no. 2, 1997, pp. 44-52.[7]J. Tyberg, "Local Adaptive Slicing for Layered Manufacturing," M.S. thesis, Mechanical Engineering, Virginia Tech, Blacksburg, Virginia, U.S.A., February 1998. URL: http://cadserv.cadlab.vt.edu/bohn/projects/tyberg/[8]1. Tyberg and J. H. B~hn,"Local Adaptive Slicing," Rapid Prototyping Journal, vol. 4, no. 3, 1998, in print.[9]Stratasys, Inc., QuickSlice Training Manual, Release 5.0, October 1997.[10]Stratasys, Inc., Utilities Manual, Release 5.0, October 1997..

M® TECHNOLOGY PROCESS IMPROVEMENTSJames W. CombWilliam R. PriedemanPatrick W. TurleyStratasys, Inc. Eden Prairie, MinnesotaABSTRACT Since the introduction of rapid prototyping technology as a tool for time compression and concurrent engineering in the design and manufacturing process, many enhancements and refinements have been made based on the experience of users and manufacturers of rapid prototyping equipment. These improvements contribute significantly to faster production of quality output from rapid prototyping systems.There are diverse control and material selection parameters that affect prototype models built using the Fused Deposition Modeling (FDM®) process. This paper reviews the role of several of these parameters in the process.

This paper reviews the role of several of these parameters in the process. Data will be presented to help the user choose the appropriate material for specific applications including density, tensile modulus, flexural modulus, tensile strength, flexural strength, impact strength, and hardness.The integration of material, hardware, and software in the FDM technology begins with the understanding of the basic requirements of the machine and ends with an operating procedure to choose the parameters for optimal model output and efficiency. Some of the variables include: part geometry, deposition geometry, deposition speed, liquefier temperature, material, flow control parameters, etc. Designed experiments are used in material formulation through modeling parameter defmition activities.INTRODUCTION The fused deposition modeling process for rapid prototyping integrates three key components of the system: software, hardware, and materials.

Each component can be analyzed independently for very simple model systems, but for real systems and broad operating conditions, the system complexity and interactions grow. The challenge to develop user-friendly FDM rapid prototyping systems is to define the relationship between the key input variables and the key modeling characteristics, or responses. Designed experiments can begin to identify the key variables and lead to strategies for optimizing the modeling process in a very efficient manner.PROCESS The FDM process forms three-dimensional objects from CAD generated solid, wire frame or surface model data through the consistent dispensing of individual layers or thermoplastics materials through a controlled temperature head. The model is built layer upon layer, from the bottom up.

The model is built layer upon layer, from the bottom up. The designed object emerges as a solid three dimensional part without the need for tooling.The process involved in the development of a three-dimensional model begins with the creation of a conceptual geometric model on a CAD workstation. The model is imported into the ProtoSlicea software program which mathematically slices the conceptual model into horizontal layers and deposition paths are created. The path data is then downloaded to the modeler. The modeler operates in the X, Y, and Z axes, basically drawing the model one layer at a time. Once the build cycle begins, no operator attendance is required.A spool of thermoplastic modeling material, .070 inches (.18 cm) in diameter, feeds into the temperature controlled FDM extruding head, heating the material to a semi-liquid state.

The semi-liquid is extruded and deposited into42A spool of thermoplastic modeling material, .070 inches (.18 cm) in diameter, feeds into the temperature controlled FDM extruding head, heating the material to a semi-liquid state. The semi-liquid is extruded and deposited into ultra-thin layers onto a fixture-less base. Once the material is directed into place by the X-Y controlled FDM head, the material solidifies, creating a precision laminate.MATERIALS Users of rapid prototyping technology tend to request models with materials whose properties are similar to materials they might use in their end-use applications, e.g. injection molding. In reality, each rapid prototyping technology rarely brings all the properties of an end-use material to the designer, and FDM is not an exception to this.

injection molding. In reality, each rapid prototyping technology rarely brings all the properties of an end-use material to the designer, and FDM is not an exception to this. Materials used in Fused Deposition Modeling must satisfy the requirements of the designer, their subsequent application, and their integration into the FDM process.As a subset of the universe of thermoplastic materials, FDM materials form strong interlayer bonds at or near their melting points and have appropriate composition and morphology to provide a relatively stress-free, low distortion part once the material has reached room temperature.An FDM material must also have an adequate flexural modulus and strength to be formed into a filament, spooled, In addition, it must have sufficiently and used as a piston to pump the material through the head, liquefier, and tip.

low viscosity to be pumped through the same hardware and also produce well-defmed road widths over a broad range of geometries and deposition rates.These requirements specify the materials which will effectively produce quality models using the FDM process. Some familiar candidate polymer chemistries include polyolefins, polyamides, and polyesters. The four materials currently available from Stratasys have the mechanical properties shown in Table 1. MWO 1 and ICW04 are wax formulations for use in the investment casting process; and P200 and P300 are plastic formulations (polyolefin and polyamide, respectively) that are stronger and have higher melt points.Thermal and rheological analysis has been performed on all of the current materials. These data have been important in designing the mechanical components of the FDM head which melt and pump the material. Figures 1-3 show this data for ICW04, the Stratasys investment casting wax.

These data have been important in designing the mechanical components of the FDM head which melt and pump the material. Figures 1-3 show this data for ICW04, the Stratasys investment casting wax. Melt viscosity is measured using a vibratory rheometer. Thermal analysis was performed using differential scanning calorimetry (DSC) and volumetric expansion by thermal mechanical analysis (TMA). ICW04 has the lowest volumetric expansion of the current materials. Table 2 lists the ash content, specific gravity, softening point (R&B), and penetration.HARDWARE Figure 4 depicts the mechanical design of the FDM head. The function of this assembly is to heat and pump the modeling material through the tip and onto the modeling·· surface to produce precise parts. It is a lightweight assembly designed to move at modeling speeds without affecting positioning accuracy.A small D.C. motor drives a set of feed wheels to provide up to 10 lbs force to push the filament through the liquefier and tip.

motor drives a set of feed wheels to provide up to 10 lbs force to push the filament through the liquefier and tip. The feed wheels, which are 1/2" in diameter and covered with an elastomer, are driven in a counter-rotating direction to provide the torque to feed the filament, which acts as a piston. The material must have sufficient colunm strength to accomplish this task; colunm strength is a function of the filament diameter, flexural modulus and strength. This provides a positive feed allowing no slippage of the filament.In the FDM process, material is deposited in layers with rectangular cross-section known as roads, and these have a width (W). and a height (Z). Coupled with the speed of the deposition (S), a volumetric flow rate which the system must provide is defined. The feed mechanism must be able to meet or exceed this flow rate over the full range of viscosities and pressure drops. The design considerations involve the material properties described earlier.

The feed mechanism must be able to meet or exceed this flow rate over the full range of viscosities and pressure drops. The design considerations involve the material properties described earlier. The pressure (P) which develops in the liquefier and tip is dependent upon the length and diameter of the liquefier ) as well as the material viscosity (h) at temperature (T), and volumetric flow rate (V). The (~,LL) and tip (ely. ,LT43tip diameter has a very large impact on the pressure drop because it is much smaller than the liquefier diameter. ForNewtonian fluids, the pressure drop can be expres.s.ed asP = 1:8hV1+~~Because pressure is force divided by area and the area is the cross-sectional area of the filament in FDM, the force is proportional to pressure as well. A force vs. volumetric flow rate curve is shown in Figure 5 for various tip sizes and liquefier combinations.

A force vs. volumetric flow rate curve is shown in Figure 5 for various tip sizes and liquefier combinations. This data was taken using the P200 material with tips of .010", .016", and .052", and liquefier temperatures of 102 and 1070C. As the theoretical equation implies, extrusion force is heavily dependent on tip size. This force is also significantly influenced by liquefier temperature, though to a lesser extent, as material viscosity decreases at higher temperatures. For a given set of operating conditions, the maximum flow rate can be determined by comparing that extrusion force with the column strength of the incoming filament.A key characteristic of a fmished model is the surface fmish which is a function of the liquefier viscosity, envelope temperature, deposition speed of the material, and the geometry of the part.

The liquefier and tip must heat the material to setpoint and the feed mechanism must pump the material out of the tip at the full range of tip diameters and material viscosities. The required heat transfer is a function of the thermal properties of the liquefier, tip, and modeling materials as well as the diameter of the filament and volumetric flow rate. Whereas it is advantageous to increase the filament diameter to increase the column strength for pumping, the reverse is true for melting the modeling material. The envelope temperature can affect the surface finish (see Figure 6) of the model by softening the material and reducing its flexural modulus. As a new layer is deposited, the previous layer may deflect downward causing preferential flow, due to a lower pressure drop and a poorly defined road width.SOFTWARE/FIRMWARE This discussion is limited to the pumping and motion control rather than the slice routines.

The software/frrmware control the motion of the head assembly on the carriage and also the motion of the feed wheels. The major tasks of feed wheel control can be broken down into two major categories: i.e., start/acceleration and stop/deceleration activities. Steady-state pumping requires very accurate carriage and feed control to assure precise geometry and road width. At the start or stop of the road, material flow is wheel inherently different and requires different motor control to accommodate visco-elastic material behavior to precisely begin or end a road.steady-state and transient behavior;Before the carriage moves to start a new road, the feed wheels meter a small amount of material in anticipation of the carriage accelerating to the steady-state. When the carriage moves, the flow is slowly turned on based on constant acceleration to the full pumping rate.

When the carriage moves, the flow is slowly turned on based on constant acceleration to the full pumping rate. Likewise, near the end of the road, the pumping rate decelerates prior to the actual end point which creates a "starving" condition at the end point. The deceleration, acceleration, and pre-start metering control values are dependent on the material visco-elastic properties at the application temperature. Each material has its own characteristics which must be programmed into the software/firmware.SYSTEM INTEGRATION The integration of material, hardware, and software in the FDM technology begins with the understanding of the basic requirements of the machine and ends with an operating procedure to choose the parameters for optimal model output and efficiency. Some of the variables include: part geometry, deposition geometry, deposition speed, liquefier temperature, envelope temperature, material, flow control parameters, etc.

Some of the variables include: part geometry, deposition geometry, deposition speed, liquefier temperature, envelope temperature, material, flow control parameters, etc. Designed experiments are used in material formulation through modeling parameter definition activities.In the development process, the sheer number of variables to be considered is overwhelming, and therefore the use of small screening designs has great utility2,3. The following example illustrates a material formulation mixture design and shows how it is organized, executed and analyzed. The purpose of this design is to screen, or identify, It is a linear design which allows analysis of individual variables only, and as a four the critical variables.44component mixture, it contains nine unique trials. Table 3 shows the design and uses coded variables to illustrate the two levels (-1 and +1) and a center point (0).Table 4 contains actual values for this example.

Table 3 shows the design and uses coded variables to illustrate the two levels (-1 and +1) and a center point (0).Table 4 contains actual values for this example. Note that the sum of the components for each trial is one. The experiment is executed by preparing the trial mixtures and measuring the responses. Table 5 shows the response matrix of some of the properties that are important to an investment casting wax. These include softening point, density, ash, viscosity and volumetric expansion. Values for all the trials are listed in this table.Next, regression analysis reveals the relationship between the variables and responses (Table 6), and the significance of each variable. The summary results figure indicates significance with stars - the more stars, the greater the significance. Finally, a graphical representation of the fit (Figure 7) can reveal meaningful trends.

The summary results figure indicates significance with stars - the more stars, the greater the significance. Finally, a graphical representation of the fit (Figure 7) can reveal meaningful trends. Inothis case, viscosity at 70 C.Another designed experiment we frequently use is the two variable, two level factorial with the center point (Table 7). An example for this type of experiment would be illustrated by Figure 6 where the liquefier temperature and air (envelope) temperature are variables, and surface finish, delamination and plugging are responses. A robust modeling zone is therefore identified. At this stage, some candidate materials may be eliminated due to lack of lamination, low flexural strength, and/or excessive viscosity.The goal of the next stage is to expand the variables constrained in the first experiment. These would include tip diameter, deposition speed, deposition geometry, part geometry, liquefier temperature, and envelope temperature.

These would include tip diameter, deposition speed, deposition geometry, part geometry, liquefier temperature, and envelope temperature. Typically, this is done over a broad range of standard models and test parts. The hard barriers previously described are defined at this stage. Material formulations can be compared to determine which of the initial formulations offered the most desirable characteristics, and can lead to additional designed experiments if the product requirements are not met.The final stage involves the definition of the flow control parameters in the software/firmware.

This isaccomplished by iteratively determining the values on test geometries within the modeling parameter envelope.CONCLUSIONS The integration of material, hardware, and software/firmware in FDM is accomplished in an efficient manner by understanding the basic independent functions of each system and using designed experiments to lead to optimal results.REFERENCESIMichaeli, W., 1984, "Extrusion Dies," Hanser, Munich.2Box, G.E.P., Hunter, W.G., and Hunter, J.S., 1978, "Statistics for Experimenters," Wiley, New York.3Wheeler, B., Betsch, R., and Donnelly, T., 1993, "ECHIP User's Guide and Reference Manual," ECHIP,Incorporated, Hockessin, Delaware.4530000010ססoo75000& 50000 E 30000 ~ s:20000v:l1ססoo8000"I 60005000300020001000 SlOO 800 700 40 10450 12260 140\' - ----70 15880 176=:::::'----90 194Data Courtesy of M. Argueso & Co.• Inc.TEMPERATUREFIGURE 1.

-Optimization and data handling for Additive Manufacturing of structural partsThomas Reiher*, Rainer Koch*C.I.K./DMRC, University of Paderborn, Pohlweg 47-49, 33098 Paderborn, GermanyAbstractAdditive Manufacturing (AM) offers high potential due to its freedom of design for structural parts. Especially in combination with FE-based topology optimization an optimal use of material and thus significant weight reductions can be expected. However, the application of AM is hampered by different additional manufacturing processes along the entire production chain and data handling induced restrictions.Disadvantages emerge from a lack of adjustment of the entire design process for AM. First the optimization algorithms are not targeted to the opportunities and restrictions of AM – represented by design rules – like the design of support structures. Secondly, the CAD software is not adjusted to AM in particular.

Secondly, the CAD software is not adjusted to AM in particular. Creating freeform shaped surfaces based on the optimization results is significantly less convenient than building defined blocks or turning parts following the needs of conventional machining. The indispensable subsequent interpretation of optimization results regarding the design rules and the possibilities of CAD- tools counteracts optimal results.This paper considers different approaches for a Topology Optimization (TO)-shape regaining on different sample parts including telecommunication satellite parts. An innovative design methodology is presented getting crucial for creating high quality designs.IntroductionEmerging requirements in weight savings for fuel efficiency or performance improvement lead to a very high demand for lightweight design of parts in a broad range of industries.

Especially in the areas of aerospace and racecar engineering weight reduction and performance enhancements are mandatory for commercial and sporting success. For space applications the weight reduction of structural parts is essential as not only the direct savings due to less propulsion have to be considered but additional payload can be transported into space, increasing the return of investment for the satellite mission.One of the most promising technologies nowadays for enabling very lightweight design is the additive manufacturing (AM) technology. AM is a tool free manufacturing process and thus enables very complex designs not possible to be manufactured by conventional manufacturing. These designs are supposed to use applied material best as unused material can be omitted without raising costs but even lowering due to less material consumption.Gaining such ‘perfect’ structures is hardly possible for designers by hand.

The designer has to be supported by automatic algorithms. One method for generating these designs is the topology optimization (TO). The topology optimization is a finite element method (FEM) based calculation method that is capable of allocating material only where it is needed. [Bend03]The combination of AM and TO carries great potential and positive impact in both directions: TO develops best shapes that are only manufacturable by AM. Though there are some hurdles to be taken in the design process. This paper details both processes, explains these hurdles and discusses different approaches of overcoming them.1092FundamentalsA broad range of manufacturing technologies is subsumed under the term ‘AM’. All of them create a part’s shape not by removing material like milling or turning, but adding material. Many of them are 2D-layer based as a complex part geometry is digitally sliced in thin layers which are than build up.

Many of them are 2D-layer based as a complex part geometry is digitally sliced in thin layers which are than build up. For high value structural parts metals processed by the selective laser melting (SLM) can be used. In this case a thin layer of metal powder is deposited and the 2D geometry in this height is molten by a laser. After the material is solidified, the next layer is deposited and molten. Thereby step by step any complex geometry can be manufactured [Geb13]. Theoretically each layer is completely independent to the foregone. Although there are some restrictions impeding this absolute freedom, it shows the high complexity available by additive manufacturing. No tools or molds are needed and nearly everything is possible even without extra costs. [HHD06]This high complexity in design enables using the material best.

No tools or molds are needed and nearly everything is possible even without extra costs. [HHD06]This high complexity in design enables using the material best. The best use of material by means of lightweight design implies a most equal stress distribution over the entire part and the use of material only where it is needed, abandoning unburdened material. Depending on the loads of a part this may come to structures with struts, sponge like material distribution or “bionic” structures. Designing these structures for complex loading conditions is nearly impossible by hand as one is not able to define the force flux analytically. The designers need to be supported by automatic working computer algorithms.Therefore the topology optimization method (TO) can be used. The topology optimization is a method based on the finite element method (FEM) for calculating the stress and strain distribution in a part based on the loading conditions.

The topology optimization is a method based on the finite element method (FEM) for calculating the stress and strain distribution in a part based on the loading conditions. The optimization algorithm interprets the results of a FE-Analysis and thereby optimizes the material allocation for using least material. Therefore the biggest available design space where material is allowed to be is defined. The bigger the space the better the result due to most direct force flux. [Har08] Topology optimization is already in use in many fields of engineering for gaining the best designs. For conventional machining these results have to be revised and adapted massively to enable the manufacturability. The resulting “bionic” structures with freeform surfaces and complex struts are hardly manufacturable and if required only with very high costs.

The resulting “bionic” structures with freeform surfaces and complex struts are hardly manufacturable and if required only with very high costs. Additive manufacturing is the key enabler for this technology as the additive manufacturing does not need any tools and the design is restricted by only very rare limitations. One of the simplest examples often overseen with its complexity for milling is the ability of producing very deep but small pockets without extra costs for special milling tools. More ambitious examples are complex undercuts due to most direct force flux following struts. Very voluminous designs leading to high moments of inertia and thus high stiffness can be realized by the combination of high burdened directly mapped struts with small, very filigree connecting bridges or thin but three dimensionally curved shear plates.Beside that very smooth and “bionic” like designs are possible.

As there is no limitation in complexity the surfaces and struts can flow in each other with very perfect and complex radii. One is not limited to the standard curve-radii but can design radii with very low or even without stress risings due to notch factors. The principle of optimizing notches as proposed by [Mat03] can be used best by the combination of TO and AM.Though some limitations still have to be considered from different sides:  process induced restrictions  conventional post processing  design tools First the design has to follow some process induced restrictions anyway. In AM for metals support structures are needed between the build platform and the actual part. These are very thin and brittle structures needed to ensure the part to stay in place, reduce distortions and1093particularly for heat conduction from part to the build platform.

These are very thin and brittle structures needed to ensure the part to stay in place, reduce distortions and1093particularly for heat conduction from part to the build platform. These are needed at the beginning under the lowest points of a part and later on under all down facing surfaces with an angle to the build platform lower than 45° or starting volumes [ZiAd14]. Furthermore neither too small nor to big material accumulations are possible. For a safe build there has to be enough material while too big material accumulations may lead to high internal stresses causing distortions up to destruction of a part [Ada15]. As well inner cavities are possible to be build, but need a connection to the outside to remove the unused powder. Overall the dimensional accuracy and tolerances of additive manufactured parts are questionable [GiWa14]. This has to be kept in mind for TO as well as for any interfaces to other parts.Secondly the conventional post processing has to be considered.

This has to be kept in mind for TO as well as for any interfaces to other parts.Secondly the conventional post processing has to be considered. After the build process the part has to be removed from the build platform and the support structure hast to be removed. This can be done mainly by hand but for series production as well as for critical surfaces machinability in these areas has to be ensured. Due to the tolerances and the quality of as-built surfaces all interfaces have to be post processed with conventional machining. This is similar to casting processes [Habo14], [Tho09]. For the combination of AM and TO one has to consider the loads during conventional machining, as the resulting shapes from TO are optimal load adapted solely for the given loads.

For the combination of AM and TO one has to consider the loads during conventional machining, as the resulting shapes from TO are optimal load adapted solely for the given loads. The resulting filigree shapes may not withstand the high loads from conventional machining and thereby have to be kept in mind either when setting up the TO-model or checked afterwards in FE-Analysis of the optimized part.A further point, especially in case of integral design, is the accessibility of tools for the post processing. If an assembly is set up by different parts it might be easier to access all relevant points before assembling. If this assembly is manufactured in one step by AM it might be critical to access mounting points for correction of the surfaces as for example for washer placement.The third point is the used software. Common design tools are not made for high complex freeform shaped 3D-parts.

Common design tools are not made for high complex freeform shaped 3D-parts. Feasible freeform software exists for art designers but not for regaining and reproducing the high complex shapes of a topology optimization. These include complex surfaces with outgrowing connectors, splitting up and growing together again. For industrial use as well parametrically designed interfaces have to be used.Topology optimization shape regaining methodologiesEven if all beforehand mentioned limitations have been considered, further manual work is needed on the TO results as they are not directly printable. The topology optimization weighs each element inside the given design space regarding its importance for the part. Thereby each element gets a density assigned from zero to one and is included correspondently in the calculation. The designer exports these results according to a chosen threshold representing feasible shapes.

The designer exports these results according to a chosen threshold representing feasible shapes. By this the results are not clear and usable for direct manufacturing as shown in Figure 1. These are three of the most common difficulties of topology optimization results:unclear results - -insufficient resolution insufficient design space1094Figure 1: Examples for shape difficulties in topology optimization resultsThe first example shows unclear results as there is a strut with a shear plane but this plane could not be clarified to struts or surfaces but could be a perfect place for AM specific low dense structures like lattice structures. The second one shows interrupted connections due to insufficient resolution. The connection bridge between two bigger areas of the part is not closed as there are some very big elements next to very small ones leading to some elements completely not shown. This result would not be useful.

This result would not be useful. The third example shows the result of insufficient design space. The inner volume of the strut is highly burdened and thereby shown red as very important. The outer shells are not that important and thereby shown green. If the exported result should have an equal stress distribution, the shown surfaces all should have the same colour and thereby importance. In this case some elements in red can be seen what is explained by missing elements in the design space. Due to the high volume of the design space the overall optimization was conducted in two steps as explained in the use case.

Due to the high volume of the design space the overall optimization was conducted in two steps as explained in the use case. On basis of the first iteration some elements where deleted and now it can be seen that some elements were deleted that would have been needed.Due to these examples and further details the designer has to revise, correct and adapt theresults [LiRei15].Especially if the part is still in design process and the exact shape of interfaces or transmission points have to get updated, it is mandatory that conventional CAD-tools can interact with the data of the optimized part. If transmission points or interface shapes depend on the attached parts or vice versa, interoperability is needed. Standard CAD-tools often are based on the kernels ACIS or Parasolid and originally are made for the use of B-REP and CSG features. Though they are already updated with freeform tools for design of NURBS, they still have some problem with organic shapes.

Though they are already updated with freeform tools for design of NURBS, they still have some problem with organic shapes. Features like NURBS, SubD or T-Splines are mathematically well described and very useful for the design of for example cars. A freeform surface organic shape part is a combination of many NURBS, each with four points and four sides. The next face has to match exactly at the other one to prevent lacks and to ensure a “watertight”, problem-free volume model [Yar13].Designing an entire structure with these functions is a time consuming and error-prone process. One example is shown in Figure 2. The target structure is not that complicated, but due to the complex handmade shapes there are useless bulges. The part has to be divided into a great amount of single surfaces, depending on the tolerated complexity of each one. The shown problems arise from freeform splines following the optimization results that do not lead to a clear defined surface.

The shown problems arise from freeform splines following the optimization results that do not lead to a clear defined surface. Especially the mandatory tangentially connection to neighboring faces may cause distortion of single surfaces as the hard radii continue and result in over interpreted following splines.1095Figure 2: Unusable bulgy structure by handmade freeform shapesDifferent approaches for designing optimal structures for AM have been explored in the past as shown in Figure 3 [LiRei15]. Therefore the use case described later on (“upright of a formula student racing car”) was already used in the respectively relevant actual configuration of current car.Figure 3: Different approaches designing optimal structures for AMThe first one is made by most direct use of freeform surfaces. This leads to a very lightweight design, as the TO results are implemented the most direct. Though, big disadvantages emerge from a very high effort and problems as shown in Figure 1.

This leads to a very lightweight design, as the TO results are implemented the most direct. Though, big disadvantages emerge from a very high effort and problems as shown in Figure 1. The second one was done without a topology optimization but based on a wire frame model, FE-analysis and conventional redesign and with regard to special direct manufacturing design rules [ZiAd14]. The results are very usable for AM but not optimal as they depend on the engineering experience and not optimization was used. The third one is based on topology optimization again but conventional CAD-features like linear extrusion were used and the design rules for reducing support and material accumulations are neglected. The fourth approach again used topology optimization and standard tools but with more lucid elements and thereby as well the TO-results as the design rules are used.

The fourth approach again used topology optimization and standard tools but with more lucid elements and thereby as well the TO-results as the design rules are used. Except the first approach rare freeform surfaces are used and thereby the topology optimization results with “bionic” elements are not represented best.A typical design procedure after TO could be a retransformation to NURBS based on all elements by complex algorithms, as rudimental already implemented by common TO- software or by designer’s interpretation as shown before. Both ways would not lead to perfect surfaces as automatic surfaces often keep bulkiness and difficulties with unclear results appearing as explained in Figure 1. A designer though may equalize result issues but is not capable of redesigning all complex details. Furthermore the manual work takes a very high effort. The first approach with most direct shape regaining took several weeks of hand work.

Furthermore the manual work takes a very high effort. The first approach with most direct shape regaining took several weeks of hand work. Therefore a more automatic and less elaborate workflow is needed.Summarizing, a new methodology for regaining TO result shapes is needed.1096Voxel based TO shape regaining methodologyOne key to solve these issues is the underlying representation methodology. Especially due to the mandatory change from B-Rep / CSG / NURBS into a polygon model for a FE- Analysis and for topology optimization this is necessary anyway. For the simulation the analytically defined and closed surfaces are meshed by a pre-processor. The result is a geometry representation based on small polygons like 2D trias or quads and 3D tetrahedrons or hexahedrons. During this transformation step information gets lost as only the position, size and shape and neighbors of each element are known.

During this transformation step information gets lost as only the position, size and shape and neighbors of each element are known. The beforehand parametrically designed structure is now fixed and all parametric dependencies are deleted. Information about defined surfaces and especially exact defined radii are no longer available. A circular hole will be presented by more or less angular elements.A proper way to keep all desired details and remove unwished details or waviness is to stay one step longer in the alternative representation level as shown in Figure 4. The polygon model is transformed into a voxel representation. This enables an even more freeform design, as the voxels can be moved, removed and added completely free. Thereby it is as easy as in no other representation to smooth surfaces, add material where needed and design surfaces perfect for stress distribution.

Thereby it is as easy as in no other representation to smooth surfaces, add material where needed and design surfaces perfect for stress distribution. Low notch factors are possible as one is not restricted to circular notch forms and may design notch forms as propagated by C. Mattheck [Mat09].Figure 4: Design Methodology based on voxel representation for regainingThe voxel representation benefits of an inherently closed volume model. The TO-results are exported in a .stl file format that only shows surface elements as typical for stl. The single stl-triangles exactly represent the mesh of the topology optimization. Thereby the stl represents fully closed volumes without surface defects. The following modifications and transformations may cause defects and problems in data. Therefore this working dataset is filled with voxel elements.

The following modifications and transformations may cause defects and problems in data. Therefore this working dataset is filled with voxel elements. Henceforth the entire part is represented by fully closed volumes and due to the voxel technique there is no possibility of producing gaps in surface with resulting needed repair work. Furthermore, voxel can be added and removed freely. Thereby it is very easy to add needed struts that may have not been exported with sufficient material,1097smooth surfaces and optimize transitions between different elements. Details can be removed by smoothing and unclear results as shown in Figure 1 can be defined either as thin walls or low dense structures like lattice structures.The voxel representation is very good in representing smooth things but it is not capable of showing sharp edges as needed for drilled holes.

Therefore in the next step two different approaches have to be considered.If a comparatively simple part is designed with only less interfaces it might be possible to circumvent this constraint and to design adequate enough surfaces with voxel technique as well. The viable sharpness depends on the chosen voxel size in terms of resolution. For bigger parts a higher resolution has to be chosen because otherwise the model is no longer processible. In this case the voxel model could be retransferred into .stl data and printed directly.If, as already mentioned above, the part shall be used further in standard CAD tools or there is a need for parametric design of interfaces or special areas of the part, the voxel model is transferred into NURBS. These are processable in standard tools.For complex optimization problems including inner cavities like hydraulic blocks this is mandatory to adjust the smoothed data as shown in Figure 5.

These are processable in standard tools.For complex optimization problems including inner cavities like hydraulic blocks this is mandatory to adjust the smoothed data as shown in Figure 5. While smoothing it is possible that the additional material needed for smoothing has left the design space and extends into areas where no material is allowed. This can be seen in step 2, 3 and 4 as the hole for mounting was completely closed during smoothing. On the other hand there might be material reduced for smoothing purposes from the non-design space. This can be seen in step 6 comparison. Any grey material can be seen is missing. At the strut there are some elements protruding that have been smoothed away as wished. Though, around the hole there is insufficient material left. Both, insufficient material and protruding material might hamper the correct working of the part and therefore is not permitted.

Though, around the hole there is insufficient material left. Both, insufficient material and protruding material might hamper the correct working of the part and therefore is not permitted. Especially for complex assemblies and application sites with non-design spaces due to moving parts a correct design space is needed to ensure a proper application.Figure 5: Workflow for combination of smoothed data with design / non-design spaceThe already existing design and non-design spaces represented by parametrically designed standard features are still available in the standard tools. These are now needed as shown in step 5.

These are now needed as shown in step 5. By use of Boolean operations on the on hand material can be added and on the other hand material can be removed to adjust and correct the smoothed shapes into usable ones.1098Use CasesTwo use cases are chosen for testing and showing of proposed voxel based TO shape regaining methodology: The first one is a structural part used in telecommunication satellites, the second one is again the already mentioned upright for a formula student race car.The considered satellite part is called “RW-Bracket” and is used to mount a reaction wheel inside the satellite. These reaction wheels can be set into rotation and thereby the satellites orientation is adjusted without using propellant. The overall structure consists of the green main bracket and three smaller blue ones as shown in Figure 6 a). This is mainly due to manufacturability and by AM the two smaller blue ones may be integrated into the Bracket.

This is mainly due to manufacturability and by AM the two smaller blue ones may be integrated into the Bracket. The third, bigger blue bracket has can’t be integrated as it is used for other mounting purposes as well.Figure 6: a) Mounted Reaction Wheel-Bracket in Satellite (without reaction wheel) b) design space c) reduced design space in second iterationFor the topology optimization first the maximum design space is defined. The bigger the design space the better the result can be as the material could be distributed best and most voluminous. Figure 6 b) shows the full FE-model including the design space in orange, the non-design space in red and the appearing forces. The non-design space is the interface towards the reaction wheel. The considered material for direct manufacturing is AlSi10Mg. Due to the size of the design space the optimization has to be conducted in two steps.

The considered material for direct manufacturing is AlSi10Mg. Due to the size of the design space the optimization has to be conducted in two steps. Figure 6 b) shows the second iteration as the design space is reduced massively based on the optimization results of the first optimization. The optimization is constraint for a maximum displacement of 0.1 mm of the center of gravity (CoG) of the RW-Bracket. This is as well the force transmission point as the relevant forces arise from the weight of the bracket during launch. During launch as well the Eigen frequencies are very critical and therefore the bracket is restricted to a minimum first Eigen frequency of 140 Hz.Figure 7: Topology optimization results of RW-Bracket1099The results with a threshold value of 0.3 are shown in Figure 7. The density and importance of elements is shown by a scale from red (very important, density one) to blue (not important, density zero). Elements with a density less than 0.3 are not shown.

The density and importance of elements is shown by a scale from red (very important, density one) to blue (not important, density zero). Elements with a density less than 0.3 are not shown. The design is mainly driven by thin legs stiffened by small bridges. It can be seen that some of the mounting points are no longer used. Therefore a special investigation was made to ensure the appearing forces at the remaining points are not too high. This shows that the additional smaller brackets are not needed from a mechanical point of view. Producing this design conventionally is nearly impossible.By use of the proposed voxel based TO-shape regain methodology the results where optimized and adapted and the structure as shown in Figure 8 a) is the final result. By help of topology optimization the weight is lowered by 60%, the displacement is lowered by 37% while the stiffness was increased as the 1st Eigen frequency is increased by 20% as shown in Figure 9.

These good results are due to the perfect load adapted design. Figure 8 b) shows a comparison of the conventional and AM design. It can be seen that the structure follows much more direct the force flux from the reaction wheel to the mounting points. In conventional design the three legs are very voluminous but end in a very small area for connection to the smaller brackets. In AM design the voluminous and thereby stiff structure is retained directly to the mounting points. This results in stiffer but lightweight structure.A further gain is the needed time for shape regaining. For comparison these result were tried to regain with standard CAD tools, even with modern freeform surface tools. Two weeks were spent and a feasible result was not attainable. Beside the very time consuming work the designed shapes were very bad and did not really fit to the optimization algorithms.On the other hand, the redesign based on voxels is realizable in about one day.

For further considerations, especially with regard to tight time schedules as for structural satellite parts, this should be considered.Figure 8: a) Final RW-Bracket Design b) comparison of conventional and AM designComparison Conventional - AM Design1200100080059 %60040020037 %+ 20 %97 %0Conv AMweight [g]1114 456Max disp [mm]0,076 0,0481st Eigen frequency 180 216B2F ratio50 1,5Figure 9: Comparison Conventional – AM Design RW-Bracket1100The second sample part again is the Formula Student racing car upright of this year’s car from the University of Paderborn1 similar to those presented in Figure 3. Though the requirements change during the years depending on the cars concept, the task is very good comparable and thereby the rough results over the years as well.Over the years the computing resources got very strong. Thereby the shown optimization was done with a very high amount of elements.

Thereby the shown optimization was done with a very high amount of elements. Hence the results are already very smooth as shown in Figure 10 a). A higher mesh resolution leads to better results especially in surface quality. Although, the described problems like gaps in connectors and unclear results will be the same with high resolution optimizations. Thereby, some details have to be defined. The result of the geometry regain is shown in Figure 10 b). It can be seen that the geometry is very close to the optimization result. With the new methodology the designer is capable to retain the structure and the mechanical principle behind the results. The outcome is a part with a very equal stress distribution over the entire structure with no critical stress risings in sharp notches as all junctions are very smooth.This result can’t be compared to a conventional design as there was no conventional design carried out.

For interpretation of the quality of the design the FE-analyzed equal stress distribution can be used. As well, again the short time for shape regaining of about one day is unachievable with conventional design. Similar conventional design of previous years had shown design times of weeks, even without freeform surfaces.Figure 10: Upright for Formula Student racecar 2015; a) topology optimization result b) shape regain with voxel methodologyConclusionAdditive manufacturing is a key enabling technology for the optimization of structural parts, topology optimization carries the potential to use this technology in a ‘perfect’ way. This combination of two strong technologies promises a higher degree of lightweight designs. Though there are some hurdles in regaining the very complex freeform surfaces calculated by topology optimization. Therefore a new methodology for voxel based topology optimization shape regaining is proposed in this paper.

Therefore a new methodology for voxel based topology optimization shape regaining is proposed in this paper. This methodology is capable of representing the complex structures most direct and thereby converting a digital material distribution into a clearly defined CAD model with the ability of including parametric elements for interfaces. The resulting structures offer interoperability with conventional CAD tools for parametric1 Formula Student is a worldwide design competition for students. Each team designs an open wheeled racecar each year. The competition is made to enhance the practical education.1101assembly design and for interconnection with conventionally manufacturing machines for post processing of interfaces. They represent the topology optimization results best, with low stress risings, an equal stress distribution and abandonment of unnecessary material.

They represent the topology optimization results best, with low stress risings, an equal stress distribution and abandonment of unnecessary material. Furthermore very high complex designs are achievable in short time and robust 3D-data can be used in the following process steps.The methodology was validated by hand of two use cases and the promised benefits can be achieved. Both use cases were set up in a short time with very satisfying results with regard to stress distribution and material savings. As well the actual additive manufactured parts show that the methodology works and it is possible to produce high complex shaped structure elements by use of additive manufacturing.For gaining optimal additively manufacturable structures there is a clear need to enhance methodologies to gain shapes with low need for support structures in manufacturing process further on. This can be done either in the optimization algorithm or in the following adaption process.

Thomas, D., The Development of Design Rules for Selective Laser Melting. University of Wales; National Centre for Product Design & Development Research, 2009. Yares, E. (2013): Organic Shape Modelling for Engineers, Design World[Tho09][Yar13] [ZiAd14] Zimmer, D.; Adam G.; “Design for Additive Manufacturing - Element transitions and aggregated structures”; CIRP Journal of Manufacturing Science and Technology; 20141103.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFEASIBILITY ANALYSIS OF UTILIZING MARAGING STEEL IN A WIRE ARC ADDITIVE PROCESS FOR HIGH-STRENGTH TOOLING APPLICATIONSChristopher Masuo*, Andrzej Nycz*, Mark W. Noakes*, Derek Vaughan*, and Niyanth Sridharan*Manufacturing Demonstration Facility, Oak Ridge National Laboratory, Knoxville, TN 37932, USAAbstractTraditional tool and die development require skilled labor, long lead time, and is highly expensive to produce. Metal Big Area Additive Manufacturing (mBAAM) is a wire-arc additive process that utilizes a metal inert gas (MIG) welding robot to print large-scale parts layer-by-layer. By using mBAAM, tooling can be manufactured rapidly with low costs. For cold work tooling applications, a high hardness level is desired to increase the life-time of the tool. A promising material that can achieve this is maraging steel.

For cold work tooling applications, a high hardness level is desired to increase the life-time of the tool. A promising material that can achieve this is maraging steel. Maraging steel is known to have good weldability; however, further testing must be conducted to ensure it is feasible for printing using mBAAM. In this paper, initial process parameters were obtained by printing single bead welds. Multi-bead walls were then printed with some refinement of process parameters to construct homogenous outer features of the walls. Lastly, the walls were heat-treated, and hardness data was gathered through Rockwell Hardness tests.IntroductionMetal Big Area Additive Manufacturing (mBAAM) utilizes a robotic welding arm toprint metal layer-by-layer, similar to Fused Filament Fabrication (FFF). Unlike traditional robotic welding, welded beads are used to form complex geometry.

Unlike traditional robotic welding, welded beads are used to form complex geometry. In mBAAM, medium- to large-scale metal objects are additively produced using a Direct Energy Deposition (DED) process. Large-scale additive systems are considered to have a print envelope of one to two meters on its longest axis [1]. The challenges that arise when printing in large-scale are uneven layer height error accumulation and warp distortion caused by heat. This relies on a closed-loop control [2] that tracks the height of the layer in real-time to ensure the printed part is near-net shape. The mBAAM system developed by Wolf Robotics and Oak Ridge National Laboratory’s (ORNL) Manufacturing Demonstration Facility (MDF) is a gas metal arc welding (GMAW) robotic system that incorporates this closed-loop control feature and has successfully printed metal parts used for tooling and other industrial applications (e.g., excavator stick, propeller).

This system is also equipped with two welding torches for multi-material metal printing. Currently, metals used for printing with this system are mild steel and 410 stainless steel. This means that tools and dies printed with the mBAAM system comprised of these metals.When developing high performance tooling for metal forming, a high hardness level isnecessary for the tool to maintain long production cycles. This is important in cold work tooling applications as the materials for forming can be particularly hard, which leads to deforming and wearing of the tool [3]. When dealing with steel sheet metal, a Hardness Rockwell C (HRC) above 50 is the minimum preferred hardness. This introduces some difficulties when using mild steel and stainless steel, as they have an HRC below 50. For printed mild steel, Shassere et al.This manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy.

For printed mild steel, Shassere et al.This manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes. The Department of Energy will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).684determined that mild steel has an HVN of 170 [4], which is below the value of HRC. Printed 410 stainless steel has an HRC of 41; however, it is at least 10 values below the recommended hardness. The advantage of mBAAM is that these tools and dies can be printed in a matter of days.

Printed 410 stainless steel has an HRC of 41; however, it is at least 10 values below the recommended hardness. The advantage of mBAAM is that these tools and dies can be printed in a matter of days. These parts can be quickly replaced; however, they would have to undergo post-process machining to get the proper surface finishes and tolerances required. Increasing the tool-life of these parts would decrease the need to produce these tools and dies, thus effectively decreasing cost.When considering an HRC above 50, machining becomes a concern, as high hardnesswould result in wearing out the machine tool quickly. In contemplation of this situation, the material should have the characteristics of a soft metal for machining but have the capability of post-process hardening. Due to the process used in the mBAAM system, materials used for printing should have weldable characteristics while in the form of a filler wire.

Due to the process used in the mBAAM system, materials used for printing should have weldable characteristics while in the form of a filler wire. For this reason, the material selected to achieve these characteristics is maraging steel.Maraging steel is known as a high-strength steel, having a tensile strength ranging from1000 MPa to 3000 MPa [5]. It also has a hardness value of HRC 52-54 [6]. These properties occur when heat treating the metal to 480-500°C for 4 to 12 hours [6, 7]. Prior to heat treatment, maraging steel is rather ductile and soft allowing it to be easily machined. These characteristics make this material highly applicable for tooling purposes. Other applications of maraging steel are seen in the aerospace industry [7], as it is used for landing gears and rocket motor cases. Maraging steel also comes in different grades of filler wire.

Other applications of maraging steel are seen in the aerospace industry [7], as it is used for landing gears and rocket motor cases. Maraging steel also comes in different grades of filler wire. MAR 250 filler wire was selected as the material used in this paper.The objective of this paper is to determine if MAR 250 can be used for mBAAM andmaintain a similar hardness level seen in pure maraging steel. The methodology used to analyze this material began with observing the weldability of MAR 250 and developing process parameters that were used to test its printability. After printing samples, they underwent a Vickers Hardness test to obtain the HRC of this material. Results were shown for weldability, printability, and hardness, and the paper concluded with future experiments and applications.MethodologyA. WeldabilityThe purpose of testing weldability is to understand the characteristics of a welded beadmade of MAR 250.

WeldabilityThe purpose of testing weldability is to understand the characteristics of a welded beadmade of MAR 250. In addition to finding the characteristics of the bead, initial welding parameters were developed and used as a baseline for the printability test. The welder equipped in the mBAAM system was an R500 Lincoln electric welder. This welder consisted of hundreds of weld modes that changed the welding waveform. For welding parameters, Lincoln Electric’s Power Mode Waveform was used for its ability to use any metal and gas type [8]. In Power Mode, welding power and pinch current can be controlled. For all waveforms, wire feed speed and welding speed (i.e., travel speed of the robot) are necessary variables for welding parameter development.In this experiment, 203.2 mm long weld beads were printed on a 203.2mm x 609.6mm x 6.35mm (l x w x t) low-carbon steel plate shown in Figure 1. Wire feed speed and pinch current were held constant.

Wire feed speed and pinch current were held constant. Weld speed, power, and shielding gas were variables that were changed throughout the experiment. The beads were grouped by the welding speed used. The shielding gases used were Ar-CO2 and Trimix (90% He-7.5% Ar-2.5% CO2).685r:========================~ySingle We ld Bea dXFigure 1. Welded bead test setupB. PrintabilityThis test consisted of printing one- and two-bead thick walls. The reason for printing walls was for simplicity and their correlation to the geometry seen in a complex part. Figure 2 shows the comparison between a wall and a more complex part in terms of print path. Walls are considered a perimeter and inset feature seen in complex parts, and thus are beneficial for the conduction of these experiments.PerimeterTwo bead thick wallMulti-material stamping tool... . . .. . . IllUUl(l{lllllllJl[ iJ~li(~~~~l([[lUWFigure 2.

. . .. . . IllUUl(l{lllllllJl[ iJ~li(~~~~l([[lUWFigure 2. Print path of a wall and a stamping toolIn the printability test, surface finish and geometric conformity were the main characteristics that were observed. Parameters that affect these characteristics are bead spacing and welding parameters (e.g., weld mode, wire feed speed, weld speed). Walls were printed with the same bead spacing of 3.0mm and a constant wire feed speed of 95.25mm/sec. Weld speed and weld mode were changed for every wall printed.C. Hardness TestVickers hardness tests were conducted on a two-bead thick wall printed in MAR 250.The wall was cut into sections and were hardness tested in three locations, as shown in Figure 3. Hardness was measured from sections with different heat treatment temperatures including one section without heat treatment. The temperatures ranged from 420(cid:4200)C to 1115(cid:4200)C. For 420(cid:4200)C to 520(cid:4200)C, isothermal temperatures were kept for three hours.

The temperatures ranged from 420(cid:4200)C to 1115(cid:4200)C. For 420(cid:4200)C to 520(cid:4200)C, isothermal temperatures were kept for three hours. For 815(cid:4200)C to 1115(cid:4200)C, the temperatures were kept for one hour and were then dropped to 420(cid:4200)C for three hours.686Three bead th ick wall\Rockwell----1..,.. __ Hardness TestLocations- - -Base PlateFigure 3. Hardness test locations on the three-bead thick wallResultsRegarding the weldability test, there were no apparent issues welding with MAR 250using the Power Mode Waveform. A wire feed speed of 95.25mm/sec was used, and a power of 3.3 kW was used. 2.5 kW was originally used for the power; however, this produced a colder weld with minimal wetting. Visible ripples in the weld and non-uniform melting of the metal droplets occurred as well.

2.5 kW was originally used for the power; however, this produced a colder weld with minimal wetting. Visible ripples in the weld and non-uniform melting of the metal droplets occurred as well. In Figure 4, it was observed that using Trimix gas improved the surface quality of the welded beads, which was caused by the additional heat added to the system. Since Trimix mainly consisted of Helium gas, it has a higher thermal conductivity than Ar-CO2 gas mixture. Different welding speeds were used, and improvements were seen using a faster weld speed of 16 in/min (6.77 mm/sec) compared to using 10 in/min (4.23 mm/sec).We ld Spee d: 16 in./min GAS: Ar-COZWe ld Speed : 10 in./min GAS: Ar-COZWeld Speed : 16 in./min GAS:TrimixIFigure 4. Single bead welds of MAR 250 with different welding speeds and shielding gas mixturesUsing the welding parameters obtained from the weldability test, a single bead wall wasprinted. This resulted in a large and rough surface finish.

This resulted in a large and rough surface finish. Figure 5 shows the printed one-bead thick wall. The sides of the wall have large uneven lumps. This feature in a single bead would not be suitable when printing multi-bead thick parts. This will lead to uneven melting causing687visible voids in the printed part. Uneven melting of the layers would occur due to the actual geometry of the bead. Since the bead itself has uneven geometry, when the beads overlap each other, the bead-to-bead spacing (Figure 6) would differentiate throughout the build. Less spacing would cause significant overbuilding throughout the layers, and more spacing would create pockets between the beads.~---,Front View~~.:.....,. ,........._ .... ~ --~ ...... ~~, _,<. ~ --.. ~,Side ViewFigure 5. Single bead wall printed in MAR 250Bead-to-bead Spacing 1 - - -1 IIFigure 6.

,........._ .... ~ --~ ...... ~~, _,<. ~ --.. ~,Side ViewFigure 5. Single bead wall printed in MAR 250Bead-to-bead Spacing 1 - - -1 IIFigure 6. Bead-to-bead spacing representationTwo-bead thick walls were printed to investigate on possible defects that could occurfrom these geometric instabilities observed in the one-bead thick wall experiment. The first set of walls (Figure 7) were printed with the welding waveform, Power Mode. It was determined that using a welding power of 1.7kW and increasing the travel speed improved the shape of the walls. Travel speed ranged from 16 in/min (6.77 mm/sec) to 30 in/min (12.7 mm/sec). Originally, 3.3 kW of power was used; however, this increased the surface roughness. The high power added excessive amounts of heat to the wall, thus causing molten metal to droop randomly during the print. In Figure 7c, the printed wall with a travel speed of 30 in/min (12.7 mm/sec) visibly has688the least surface roughness.

In Figure 7c, the printed wall with a travel speed of 30 in/min (12.7 mm/sec) visibly has688the least surface roughness. Although there were improvements, the printed walls still encountered uneven surface geometry.Lumpy surface (Weld speed: 16 in./min)c.Highest weld ing speed: 30 in./minBead spacing too wideFigure 7. Two-bead thick walls printed with Power Mode. a.) front view of the walls, b.) side view of the walls, c.) wall with fastest travel speed used for this set.A second set of two-bead thick walls (Figure 8) were printed with the welding waveform, surface tension transfer (STT). In STT, current automatically adjusts the heat without depending on the wire feed speed [9]. This minimizes the heat added to the weld, which is ideal for this material. In addition to minimizing heat, STT has more variables to control such as peak current, background current, and tail-out speed. This allows finer tuning of the welding parameters.

In addition to minimizing heat, STT has more variables to control such as peak current, background current, and tail-out speed. This allows finer tuning of the welding parameters. The outcome of these walls was significantly better in surface roughness. It was evident by visual observation and was confirmed by software developed in MDF. This software (Figure 9) uses scan data to interpret surface roughness by cross-sectioning the point-cloud of a wall and analyzing each side of it. The surface roughness average data is shown in Table 1. Based on the results in Table 1, walls printed with STT have about a 30% lower surface roughness average compared to walls printed with Power Mode. The walls printed with STT had higher consistent bead geometry with less waviness at the top surface of each wall.

The walls printed with STT had higher consistent bead geometry with less waviness at the top surface of each wall. There was no significant difference in wall surface roughness when increasing the weld speed from 24 in/min (10.16 mm/sec) to 30 in/min (12.7 mm/sec).Fastest: 30 in./minSlowest: 24 in./minFigure 8. Two-bead thick walls printed with STT689~ fom,lfll/'tlbqnNRl,jS•U21 lollbqnNRMS•lJ.ruFigure 9. In-house software to find surface roughness average from point-cloud dataTable 1. Roughness averages of the two-bead thick walls printed with Power Mode and STTPower ModeCross Sections Right Ra (mm)1 2 3 4 5 Average Ra0.279 0.401 0.381 0.312 0.309 0.3364Left Ra (mm) 0.383 0.319 0.419 0.344 0.319 0.3568STT Right Ra (mm) Left Ra (mm)0.303 0.252 0.202 0.189 0.222 0.23360.213 0.266 0.198 0.258 0.346 0.2562In Figure 10, hardness values are shown for each heat treatment temperature used.

As expected from maraging steel, the HRC values for no heat treatment were below 32. For heat treatment temperatures ranging from 420(cid:4200)C to 520(cid:4200)C, hardness was higher near the bottom of the wall. However, for 815(cid:4200)C and 1115(cid:4200)C, hardness was at its highest in the center of the wall. This was most likely caused by the weld dilution caused by welding on the base plate for the bottom portion. This area would be less pure in maraging steel as it has mixed some composition with the base plate. In the top layers of the wall, hardness was at its lowest. This was also seen in mild steel [4], which was caused by the less refined microstructures seen in this region.15mmNo Heat Treatment(420"C-3 Hrs l 'HSC8HRC!,480"C-3 Hrs15mm15mm815°C-1Hr+ 420"C-3 Hrs Figure 10.

Hardness values taken from three locations of a two-bead thick wall as well as different heat treatment temperatures520"C-3 Hrs1115°C-1Hr+ 420"C-3 Hrs690ConclusionMaraging steel shows promise in providing the desired hardness of an HRC above 50. Byusing MAR 250 welding wire, weldability was tested and shown to be weldable. Different shielding gases were tested with Trimix gas producing the best results. Using the starting welding parameters from the weldability test, walls were printed to observe how the material bonds together. Results show that one-bead and two-bead thick walls were successfully printed; however, the geometric features of the walls were rough, with large waviness. Using another welding mode, STT, surface roughness of the beads improved with more homogenous layers.Overall, process parameters for MAR 250 will need further adjustments to produce similar results shown in printing mild and stainless steel.

On the other hand, its printability success makes it a feasible material for mBAAM. MAR 250 did result in a slightly lower hardness than the desired 50 HRC; however, it is significantly better compared to mild and stainless steel. Future tests will be conducted with MAR 350 welding wire to observe if the hardness has increased with the higher strength material. Ultimately, maraging steel will be used for multi-material prints (Figure 11) to develop high strength tools and dies with low costs. This can be achieved by printing an outer perimeter of maraging steel and an inner core of mild steel.PrintedMachinedMild steel fill material410 stainless steel outer surfaceFigure 11. Multi-material stamping tool with an outer surface of stainless steel and an inner core of mild steelAcknowledgementsThis material is based upon work supported by the U.S.

bricating Metal Parts from 17·4PH StainlessSteel PowderGuohua Wu*, Noshir Langrana*, Sriram Rangarajan**, Rajendra Sadangi **, Amhad Safari** Stephen C. Danforth**of Mechanical and Aerospace Engineering, Rutgers,TheUniversity of New Jersey, 98 Brett Road, Piscataway, NJ 08854.** Department of Ceramic and Materials Engineering, Rutgers, The State University of New Jersey, Piscataway, NJ 08854.Abstractstainless steel is known to provide an attractive combination of high strength and corrosion resistance. In this research, the feasibility of SFF fabrication of high density parts using PH powder is examined. A part can be fabricated using both indirect and direct methods.The indirect method includes making a negative RTV mold, making compounded material using ECG binder and stearic acid with the metal powder, and pouring the compounded material to get a green part. This is followed by binder bum out(BBO) and sintering cycles.

This is followed by binder bum out(BBO) and sintering cycles. The direct method uses Fused Deposition of Metals(FDMet). In FDMet, the 17-4PH powder is compounded with a binder and extruded into filaments, followed by part building, BBO and sintering. The initialresults of the indirect method of fabrication produced 91 % theoretical density of 17-4 PH parts with Vickers hardness of 223.Keywords: Rapid Prototyping RTV mold17·4PH FDMet Binder Burning Out(BBO) SinteringIntroductionSolid Freeform Fabrication emerged about ten years ago, due to the advantage of makingcomplex parts with low cost and high speed. It is becoming one of the most important new techniques in the manufacturing area. Although there are many advantages, there exists some drawbacks. For example, in many cases, a part can be used as a physical prototype, but can not withstand high service stresses and temperature.

Although there are many advantages, there exists some drawbacks. For example, in many cases, a part can be used as a physical prototype, but can not withstand high service stresses and temperature. Various efforts have been focused on developing new process to make high strength, low cost parts with SFF, such as DTM Rapid Tooling[1], MIT 3DP[2], and Fused Deposition of Ceramic[3].The objective of this study is to develop an efficient method to fabricate metal partsquickly with low cost. Two methods are adopted in the research, one is the direct method or . Fused Deposition of Metal(FDMet), and the other is RTV Molding method. Building a metal part by FDMet is our final goal, the indirect method is used as an initial step to determine the right material composition.

Building a metal part by FDMet is our final goal, the indirect method is used as an initial step to determine the right material composition. In this paper, full details of this RTV Molding technique are reported.479Experiment ProcedureTwo different approaches are under investigation for manufacturing functional metalcomponents referred to as the "direct" and the "indirect" techniques as Fig. 1. The direct method (Fused deposition of metals, FDMet) uses metal particle loaded thermoplastic filament as the material feedstock, fed via counter rotating rollers into a heated extruder, referred to as a liquefier. The filament acts as the piston to extrude the molten metal loaded polymer material out of a 250 micron to 635 micron diameter nozzle onto a z-stage platform, where the material cools rapidly and bonds to adjacent layers. The material deposition rate and liquefier x-y position, etc., are controlled by the computer.

The material deposition rate and liquefier x-y position, etc., are controlled by the computer. For powder metal parts, the binder is subsequently removed in a thermally driven process called binder bum out (BBO), then sintered to full density. In the indirect method or RTV Molding method, a STL file is generated from the desired part, the file is delivered to the SLA or FDM machine which builds a positive prototype. The RTV mold is made by pouring a molding material around the prototype. This is done by mixing the metal powder with the ECG2 and stearic acid, pushing the mixture into the RTV mold. The binder removal and sintering process of the indirect method are the same as those of the FDMet.The development of the direct and indirect SFM processes for hard tooling are going onsimultaneously. Several steps in each of these processes have been accomplished.

Several steps in each of these processes have been accomplished. The processing details and the preliminary results are described below.Hard ToolingCompounding of Metal Powder and BinderDirect SFM ProcessIndirect SFM ProcessFilament FabricationSFM of the Tool Positive and RTV Mold FabricationFused Deposition of MetalsPour the Compounded Material into the MoldFig!. Flow chat ofFDMet and RTV mold methodsMaterialsThe material used in this research was 60% volume 17-4 PH stainless steele-10 micronswater atomized), provided by Ametek company in PA of U.S.A. The powder part~cles were irregular in shape, the chemical composition is Fe-0.031C-4.39Ni-15.82Cr-0.12S1-0.21Mn-4803.78Cu-0.014S-0.22Cb-0.610-0.02P-0.09N, and its particle size distribution (as provided by the manufacturer) is given in the Table 1.Table 1.

The size distribution of the -10 microns 17-4PH SS powderMicrons 62 44 31 22 16 11 5.5 3.9Percent(less than) 100.0 100.0 96.2 81.7 59.6 37.9 10.3 3.0The binder used in this research was ECG2(Electro Ceramics Group binder #2),which wasdeveloped by the Ceramic and Materials Engineering Department of Rutgers University[4].Stearic acid is a common surfactant for metal powder, although there are some othersurfactants that could be used such as zinc stearate or magnesium stearate.Feedstock preparationIn order to get full dispersion of the stearic acid, ball milling was used. First some grindingmedia and a certain volume percent of metal powder were added into a bottle, with some stearic acid added to acetone.

First some grindingmedia and a certain volume percent of metal powder were added into a bottle, with some stearic acid added to acetone. The mixture was full dissolved and then added to the bottle, after the mixture was ball milled about 20 hrs, it was then passed through two sieves, one to remove the grinding media and one to sieve the powder, followed by drying.The binder and coated powder are mixed using a HAAKE torque rheometer at a temperature of 165°C. First, all of the ECG binder is added to the hot mixer and melted. Next, the stearic acid coated metal powder is added into the mixer in four sub-steps (40%, 30%, 20% and 10% of the total amount of metal powder) at 15 minute intervals. After all of the metal powder is added, and when the torque in the mixer is constant, i.e. reaches a steady state value, the temperature was lowered to 130°C, after another hour a uniform compound is achieved. The whole process usually takes three to four hours.

reaches a steady state value, the temperature was lowered to 130°C, after another hour a uniform compound is achieved. The whole process usually takes three to four hours. After the compounding is completed, the compounded material is removed from the mixer, and grounded into small pieces.This granulated, compounded material is now ready either for RTV molding of the part inthe indirect method, or for fabrication of filament (by extrusion) for direct part fabrication by FDMet.RTV mold makingRTV tooling is a process which creates a soft rubber mold from an original positiveprototype structure. In this work, the desired positive tool is made by stereolithography. The481silicon material of the soft mold can withstand the 100-200°C temperature of the metal filled binder formulation that will be cast into the mold. In addition, the fabrication process is fast, accurate and inexpensive.

In addition, the fabrication process is fast, accurate and inexpensive. The RTV mold process used in this project is as follows: (1) Prepare the base, curing agent and a container. (2) Stick wax on the bottom of the original positive part so that the part can be attached to thebox, and prevent creation of air bubbles.(3) Record the weight of the base mold material (approximately 40 g), use the curing agent (10% by weight of the base), stir the base and agent, and then put them into the vacuum oven to remove any air, pour the material around the part, making sure that the part is completely covered.(4) Put the container (which now holds the uncured mold material and part) into the vacuumoven again. This is performed to remove any air trapped in the mold material.

This is performed to remove any air trapped in the mold material. Continue this process until all of the air bubbles are removed.(5) The mold material is fully cured by keeping it at room temperature and 1 atm pressure for 24hrs, then the positive part is removed from the cured mold material.After the RTV mold is ready, the compounded material is poured into the mold to form agreen part.De-binding and SinteringBefore sintering, the green part must be processed through a treatment to extract the organic binder. Based on earlier work, a heat treatment approach was developed which has successfully removed the binder from the green 17-4PH metal part. The development of an appropriate heat treatment to bum out the binder requires that we have a good understanding of weight loss and shrinkage characteristics of the material system during heating.

The development of an appropriate heat treatment to bum out the binder requires that we have a good understanding of weight loss and shrinkage characteristics of the material system during heating. Stresses can develop in the parts during heating if the binder decomposition gases evolve in the interior of the component without a sufficiently easy diffusion path to the surface. During the bum out process, the green part was buried in a very high surface area carbon powder to aid in the removal of the binder through wicking (via capillary suction) at low temperature. An optimized binder burnout schedule will be determined by measuring the shrinkage and weight loss after a finite hold at various temperatures, followed by furnace cool to room temperature.The BBO schedule for hard tooling was derived from the current PZT project [4]. It is a slow process. Following the BBO process, the part is sintered in hydrogen.

It is a slow process. Following the BBO process, the part is sintered in hydrogen. The sintering can be performed at temperatures as low as 112l oC, but higher temperatures will yield better mechanical properties. The cooling process can not be slower than what it would be in still air, and the part must be cooled to below 32°C prior to precipitation hardening heat treatment process[5].Table 3. Sintering schedule for 17-4 PH metal partsTemp(OC) 20-1260 1260 1260-20Rate 5°C/min. hold 30min. Natural coolingTime (Hrs) 4.1333 0.5 - 5482IJAA'''''' .........

Sintering schedule for 17-4 PH metal partsTemp(OC) 20-1260 1260 1260-20Rate 5°C/min. hold 30min. Natural coolingTime (Hrs) 4.1333 0.5 - 5482IJAA'''''' ......... VA.H.J were investigateda this process was improvedpartaa more complex lugpart.part wasto demonstrateprocess.was exported to the Maestrogenerated data files werewas generated of Stereolithography controlto run thewrench part.mold was madeapplying was made by A.UUAAA"" process took a long time and sintering process wasmolding process onas into mechanical properties thegreen wrench asthecompoundedwrench mold.wrench part were part into a Alumina tube at high .....H ...

JllJ'-'Adone by puttingas Table 3.reason the easilypartwas broken is that thepart was soas tobroken.FigA Green wrench partSintered Wrench part, broken during theofsinteredwere performed.hardness of the sintered metal is B97 orVickers hardness ofThe microstructure consists of ferrite grains with ASTM grain size of11 asIt was also found that the sintered part was notdense asThe porosity is483approximately 8%-9%. The distribution of the Pore Size is as Table 4.

The distribution of the Pore Size is as Table 4. The Pore Size Range is 0.5-5.5 microns.36Fig.6 Microstructure of sintered part.Fig.7 Pore size graph of the sintered part.Table 4 Pore Size Distribution of the Sintered Wrench PartSize, Microns Less than 1.0 1.0-5.0 More than 5.0Distribution -#/Inch23270 1048 635Fabrication of the lugpartSince the wrench was not a complex geometry, further efforts were made to fabricate amore complex part, the lug fit part which was required by the research program.Since the lug fit part is hollow inside, care must be taken during the RTV mold makingprocess. Since the process was done in a vacuum, if the SLA part was not attached to the bottom of the container firmly, the SLA part would rise in the RTV material, and RTV material would fill inside the lug fit part. A de-molding agent was also preferred because it was very hard to remove the SLA lug fit part from the mold.

A de-molding agent was also preferred because it was very hard to remove the SLA lug fit part from the mold. The mold was cut in half, and the SLA lug fit was removed. A piston RTV positive mold was also made to push the compounded material into the mold. The SLA lug fit, Green part, BBO part and sintered part can be seen in Figures 8 to 11. The dimension of the part in various stages are shown in Table 5.Fig.8 SLA lug fit partFig.9 Green lug fit part484Fig.l0BBO lug fit partFig.ll Sintered lug fit partSince the lug fit part was hollow inside, additional supports were added when it was putinto the sintering oven. Initially, no additional support was used and the part collapsed during the sintering, after correcting the problem, a successful Lug fit part was fabricated as shown in Fig.ll .Shrinkage of the green to sintered part was measured. The results are shown in Table 5. The maximum shrinkage of 8.64% in diameter in the sintered part was found.

The results are shown in Table 5. The maximum shrinkage of 8.64% in diameter in the sintered part was found. The shrinkages in all three dimension were similar. The shrinkage effect on the final dimension of the part can now readily be incorporated into the whole design process to gain higher accuracy.Table 5 Shrinkage Measurement ResultsSLA Inner Diameter(in) 0.466 Middle Outside Diameter(in.) 1.053 Bottom Outside Diameter(in.) 1.353 Height(in.) 0.935 * The data in the '0' is the increase/decrease percent with respect to the SLA PART.GREEN PART * 0.471(+1.07%) 1.059(+0.57%) 1.368(+1.1 %) 0.974(+4.17%)SINTERED PART* 0.430(-7.7%) 0.962(-8.64%) 1.245(-7.98) 0.880(-5.8%)ConclusionsThe present experiment work showed that using 17-4PH metal powder to fabricate a partdirectly and indirectly is very promising.

From the results, the following conclusions can be drawn as follows: (1) Materials for the hard tooling formulation were investigated, SS 17-4PH with ECG2binder and stearic acid as a surfactant.(2) RTV mold development for 17-4PH SS powder has been demonstrated. Several prototypehard tool structures have been fabricated and are being evaluated.(3) Compounding of SS 17-4PH with ECG2 has been performed and the investigation on thefilament fabrication is underway.AcknowledgmentsThis research reported here has been performed towards the partial fulfillment of therequirements for a thesis of the first author. The authors would like to thank the New Jersey485Commission of Science and Technology for the financial support under Grant #97-2890-14 through the NJIT Multi-Lifecycle Engineering Research Center Program. Also the authors wouldto thank Drs. Reggie J.

10-12,1998. [5] John Steel Processed by Conventional Powder Metallurgy techniques", International Conference and Exhibition on Powder Metallurgy and Particulate Materials, May 8-11, 1994, Toronto, Canada.University of Texas at Austin, Texas, Aug. 1 13,1997. Clancy, V.Jamalabad "Fused Deposition of Ceramics: Progress Towards a Robust andReinshagen, John C. Witsberger," Properties of Precipitation Hardening Stainless486.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFEASIBILITY STUDY OF LARGE-FORMAT, FREEFORM 3D PRINTING FOR ON-ORBIT ADDITIVE MANUFACTURINGD. Jonckers*, O. Tauscher*, E. Stoll†, and A. Thakur*‡ *Institute of Space Systems, Technical University Braunschweig, Braunschweig, Germany †Institute for Space Technology, Technical University Berlin, Berlin, Germany ‡Corresponding author: aditya.thakur@tu-braunschweig.deAbstractLarge scale, on-orbit additive manufacturing (AM) and assembly is being considered as a modular and resource saving approach to facilitate permanent human presence in space. To realise this, a novel AM approach to freeform fabricate large, functional structures in space has been developed.

To realise this, a novel AM approach to freeform fabricate large, functional structures in space has been developed. Combining the reach of a free-ﬂying CubeSat with a collaborative robotic arm and a 3D printer, large support-free thermoplastic structures can be manufactured beyond the size of the setup itself. The feasibility of the proposed fabrication approach was established using the Experimental Lab for Proximity Operations and Space Situational Awareness (ELISSA) system, where a modiﬁed fused ﬁlament fabrication setup was mounted on a free-ﬂyer to 3D print free-standing structures.

Using a continuous navigation path incorporating an inﬁnite fabrication loop, over 70 centimetre long, support-free trusses were produced to well demonstrate the potential of the proposed method in boundless direct printing of complex structures, independent of gravity or printing orientation.Nomenclature and AbbreviationsAM DOF ELISSA Experimental Lab for Proximity Operations and Space Situational Awareness FFF GNC ISM ISS JWST OBC PEEK PID PLA ROS SMCAdditive Manufacturing Degrees of FreedomFused Filament Fabrication Guidance, Naviation and Control In Space Manufacturing International Space Station James Webb Space Telescope On-Board Computer Polyether Ether Ketone Proportional-Integral-Derivative (Controller) Polylactic Acid Robot Operating System Sliding Mode Controller15081. IntroductionManufacturing large, permanent human settlements in space was popularised in the 1970sby American physicist Gerard K.

IntroductionManufacturing large, permanent human settlements in space was popularised in the 1970sby American physicist Gerard K. O’Neill, who proposed cylindrical habitats, each capable of supporting a population of over 10 000. Launching such large structures in to orbit from Earth would be impractical, and as such O’Neill proposed utilising raw material from the Moon, and manufacturing these habitats in situ [1]. In more recent years, Dunn et al. and Moraguez et al. highlighted the advantages of free-form fabrication processes such as 3D printing for In-Space Manufacturing (ISM) [2, 3]. While on demand and spare part production play a contemporary role in the operation of space stations, both authors emphasize the importance of large structure ISM surpassing the limits of deployable structures.

For example, space-based optical telescopes such as the yet to be launched James Webb Space Telescope (JWST) must be designed to stow within a volume dictated by a launch vehicle’s fairing [4]. This results in the spacecraft needing to undergo numerous deployments, including the release of 178 mechanisms to reach its operational state [5]. As deployment mechanisms are a known source of spacecraft failure, reducing the need for these could result in more reliable spacecraft [6]. Another advantage of in-space manufactured parts is the neglection of mechanical loads during the spacecraft launch – in the case of the JWST of up to 4.55 g1.

Another advantage of in-space manufactured parts is the neglection of mechanical loads during the spacecraft launch – in the case of the JWST of up to 4.55 g1. Eliminating these loads can lead to structures that are optimised for its in-space application rather than launch, in turn leading to lighter structures and reduced launch costs [7].Research into ISM techniques for large structural components has been conducted at least since the 1970s, when General Dynamics, under contract from NASA, produced a prototype “beam builder” capable of producing hundred meter scale truss segments on orbit [8]. This machine used ultrasonic welding to manufacture a beam from single-ply graphite/glass impregnated with a polysulfone resin, and though the feasibility of the concept was proved, did not ﬂy.

This machine used ultrasonic welding to manufacture a beam from single-ply graphite/glass impregnated with a polysulfone resin, and though the feasibility of the concept was proved, did not ﬂy. No investigation was done as to how more complex structures could be manufactured.More recently, the focus has shifted to additive manufacturing processes with Tethers Un- limited, also under the contract from NASA, proposing a concept termed SpiderFab™ [9]. This robot spacecraft would use an extruding nozzle, or spinneret, to extrude Carbon ﬁbre reinforced polyetheretherketone (PEEK) segments. Its robotic arms would then, in conjunction with a joining tool, assemble these segments into large structural components. A prototype truss manufacturing machine was also built, using Twintex® , a continuous ﬁbre reinforced thermoplastic yarn, to produce triangular trusses several metres in length. These could then be manipulated by SpiderFab™ to assemble higher order truss structures.

These could then be manipulated by SpiderFab™ to assemble higher order truss structures. A concept for a “trusselator” machine which could man- ufacture triangular trusses from rolls of Carbon Fibre /PEEK composite tape was also designed. The design ﬁt into a 3U (30 cm × 10 cm × 10 cm) volume, and could therefore be integrated into a CubeSat class spacecraft. In a later study, a prototype of this concept was built, though was not tested in a space representative environment [10]. A concept similar to SpiderFab™ named Archinaut has been proposed by Made In Space Inc., which would use a combination of a 3D printhead, moving in two dimensions, and a robotic traverse system moving in a third, to allow the manufacture of 3D parts. One or more robotic manipulators would allow the assembly of printed parts into larger structures [11].

One or more robotic manipulators would allow the assembly of printed parts into larger structures [11]. Again, truss elements are identiﬁed as a key application for ISM,1Ariane 5 User’s Manual Issue 5 Revision 3, https://www.arianespace.com/wp-content/uploads/2016/10/Ariane5-users-manual-Jun2020.pdf1509with uses such as supporting solar sails, or forming the backbone of future space stations suggested. A prototype of this AM mechanism was produced which successfully printed a structural element in a space-like thermal-vacuum environment. Additionally, in a shirt-sleeve environment, a beam was manufactured with a length of 37.7 m, illustrating the feasibility of manufacturing structures larger than the machine itself [12]. Both concepts share the idea of separate fabrication and assembly processes.

Both concepts share the idea of separate fabrication and assembly processes. Whilst this allows for the manufacture of complex structures in 3D space, it increases the complexity of operations and therefore reduces the robustness of the system.Fused Filament Fabrication (FFF) 3D printers have been used in space since 2014, when the 3D Print experiment was operated on the the International Space Station (ISS) by Made in Space Inc., proving the feasibility of the method in microgravity [7]. In 2016, a second FFF 3D printer was sent to the ISS, again by Made in Space [13]. The main stated aim of these printers is to allow the manufacture of replacement parts for crewed spacecraft, thereby reducing the logistical challenge of long duration human spaceﬂight [14].

The main stated aim of these printers is to allow the manufacture of replacement parts for crewed spacecraft, thereby reducing the logistical challenge of long duration human spaceﬂight [14]. As such, whilst these printers are capable of manufacturing complex components, these must be smaller than the printer dimensions, and are therefore not suitable for large scale ISM.In microgravity, trusses could be manufactured by simply extruding material for the longerons and struts into 3D space. On Earth, support-free printing of three-dimensional wireframe parts using conventional FFF 3D printers has been investigated as a method for reducing the time required to manufacture prototype structures. It was found that to extrude material directly into 3D space, it is required to keep the extruded material under tension whilst cooling it to its solidiﬁcation temperature.

It was found that to extrude material directly into 3D space, it is required to keep the extruded material under tension whilst cooling it to its solidiﬁcation temperature. The Wireprint study achieved this by pausing at the top of each upward stroke, the material would then cool before the downward stroke was printed. It was also found that the geometry of support-free structures are restricted by the nozzle geometry, in this case the authors were limited to a height of 6 mm between layers, and an angle of no more than 32° with the horizontal for downward strokes [15].In this paper, we suggest a concept which combines a free-ﬂying satellite with a robotic arm, itself equipped with a 3D printhead end eﬀector. This would combine the ﬂexibility of a 3D printer with the reach of a robotic arm to enable the manufacture of large structures.

This would combine the ﬂexibility of a 3D printer with the reach of a robotic arm to enable the manufacture of large structures. Support-free 3D printing could be achieved by inducing the spatial movement of the printhead through the combination of the satellite’s propulsion system and the robotic arm. The proposed concept has an advantage over “Archinaut” and “Trusselator” due to its ability to manufacture in all three dimensions rather than a single axis. Target applications could also include the repair of existing structures. Furthermore, it has been shown that 3D printed thermoplastic structures can be recycled back into ﬁlament, which can then be used print new structures [16]. This results in more sustainable manufacture, and could be used to alter the size and shape of the spacecraft throughout its lifetime.

This results in more sustainable manufacture, and could be used to alter the size and shape of the spacecraft throughout its lifetime. For example, at the end of its life, the cross-sectional area of the spacecraft could be increased to reduce the time until re-entry.This ﬂexibility comes at a cost however, as it requires sophisticated control algorithms for synchronised movements of the satellite and robotic arm. This paper presents a feasibility study of using a robotic arm on a small spacecraft to allow unlimited part manufacture in all three dimensions. A free-ﬂying satellite mock-up is used on an air bearing table to simulate the friction-less environment of space, in conjunction with an industrial robot arm and a 3D printhead for fused ﬁlament fabrication.1510Two methods of printing large structures – continuous movement of the free-ﬂying satellite mock-up and robotic arm, and a segmented joining of individual trusses – are tested.2. Experimental Setup2.1.

Experimental Setup2.1. Satellite Simulation Test BedThe design and test of satellite Guidance Navigation and Control (GNC) algorithms requires the replication or simulation of the microgravity space environment. These can be tested on Earth by using the Experimental Lab for Proximity Operations and Space Situational Awareness (ELISSA). This is a 4 m × 7 m active air bearing table consisting of several panels at the Institute of Space Systems at TU Braunschweig [17]. So called “free-ﬂyers” are fan-propelled satellite mock-ups that ﬂoat on an air cushion which is generated by a blower system located underneath the table. The free-ﬂyers experience friction-less movement in three degrees of freedom – 𝑥, 𝑦 and 𝜃 (rotation about 𝑧), allowing spacecraft GNC algorithms to be tested in a representative environment.

The free-ﬂyers experience friction-less movement in three degrees of freedom – 𝑥, 𝑦 and 𝜃 (rotation about 𝑧), allowing spacecraft GNC algorithms to be tested in a representative environment. It has been found that algorithms proven to work in 3 degrees of freedom (DOF) can be readily extended to the 6 DOF experience in space [18, 19]; therefore, by using a free-ﬂyer equipped with a robotic arm and 3D printer, the feasibility of using currently available spacecraft GNC algorithms for manufacturing large structures can be assessed.The free-ﬂyer used for the experiments described in this paper is shown in Figure 1a and consists of three modules – a propulsion module, containing eight fans; a service module, containing the on-board computer (OBC), as well as a battery; and the payload module, which supports a robotic arm, as well as an Ultimaker 2 extruder and printhead2.

The fans and 3D printing equipment are controlled by microcontrollers; the OBC communicates with these via a serial USB connection. The six-axis industrial Meca500 robotic arm3 has a reach of 260 mm and contains an integrated controller, this communicates with the OBC via Ethernet. The free-ﬂyer represents a small spacecraft with an accompanying robotic arm, in this case an oﬀ-the-shelf industrial robotic arm is used due to its proven performance. In the future, a robotic arm which is optimised for the task of ISM could be investigated, which may have more or less degrees of freedom.Robot Operating System (ROS) [20] is used to facilitate communication between the free-ﬂyer and a ground station, that is a Linux PC which runs the GNC and 3D printing algorithms. Finally, an OptiTrack4 optical tracking system is used to measure the position of the free-ﬂyer. This data is shared in real time with the ground station via Ethernet.

Finally, an OptiTrack4 optical tracking system is used to measure the position of the free-ﬂyer. This data is shared in real time with the ground station via Ethernet. Figure 1b illustrates how the subsystems are interconnected.A raised glass platform with paper base is used as the printing substrate, allowing the robotic arm to reach, and was found to provide good adhesion of the ﬁrst printed layer. Though the robotic arm has six degrees of freedom, the nozzle was kept perpendicular to the print surface, as with a standard Cartesian printer. Future experiments will investigate the impact of varying the angle of the2Ultimaker, The Netherlands 3Mecademic, Canada 4OptiTrack, USA1511printhead to vertical.

Future experiments will investigate the impact of varying the angle of the2Ultimaker, The Netherlands 3Mecademic, Canada 4OptiTrack, USA1511printhead to vertical. Polylactic Acid (PLA) ﬁlament5 with a diameter of 2.85 mm was used due to its good printability characteristics, as well as low melting temperature, and sustainable manufacture [21].Propulsion ModuleService ModulePayload ModuleRobotic ArmFilament Roll,ExtruderMechanismand MicrocontrollerPrintheadFree-flyerPayload ModuleService ModulePropulsion ModuleOBCEthernetUSB SerialWiFiRobot ArmOptiTrackMicrocontroller3D PrinterGround station(a)(b)Fig. 1 (a) Subsystems of the free-ﬂyer with side and rear view, (b) a schematic showing the architecture of the free-ﬂyer and how it interacts with the ground station and the OptiTrack system5Anthrazit V2 - Das Filament, Germany1512123(a)(b)Fig.

2 Two diﬀerent approaches of printing a long truss structure (free-ﬂyer movement depicted by red arrow and print head movement depicted by green arrow). (a) Shows the continuous printing approach with the free-ﬂyer moving along 𝑥-axis and the printhead on the robotic arm end eﬀector moving in the 𝑦-direction. (b) Shows the segmented printing approach with free-ﬂyer only moving in the positive 𝑥-direction and maintaining its position while printing each segment. The printhead end eﬀector moves along both the 𝑥- and the 𝑦-axis2.2. Free-Form PrintingTwo methods for manufacturing large structures were investigated. The ﬁrst combined the movement of the free-ﬂyer with simultaneous extrusion (shown in Figure 2a), this would allow the creation of structures as large as the air bearing table, or in space, as large as the distance the spacecraft could travel. The shape of the structure was dictated by the trajectory of the free-ﬂyer, along with the movements of the robotic arm.

The shape of the structure was dictated by the trajectory of the free-ﬂyer, along with the movements of the robotic arm. For traditional layer by layer printing, to ensure a consistent print, the extrusion velocity is recommended to match the velocity of the printhead in 3D space, which is a combination of the free-ﬂyer and robot end eﬀector velocities. The position of the free-ﬂyer also needs to be tightly controlled to ensure that the printed structure matches the design. The segmented method splits a larger structure into smaller segments. Each segment would then be printed whilst the free-ﬂyer held a ﬁxed position (shown in Figure 2b). By overlapping the segments, a single large structure could be manufactured. Though simpler from a control perspective, this introduces joints into the large structure, which may reduce overall structural strength.

Though simpler from a control perspective, this introduces joints into the large structure, which may reduce overall structural strength. Additionally, any error in the position of the free-ﬂyer could impact the shape of the printed structure.Structures printed in this study are support-free truss structures, as these have been identiﬁed as one of the most useful candidates for large scale ISM [8, 9, 10, 12]. As described in Section 1, the height between layers, and angle of downward strokes, are restricted by the nozzle geometry, in our case we were limited to a height of 5.5 mm and an angle of less than 45° with the horizontal. A printing temperature of 200 °C was used; though this is on the lower end of temperature range1513typically recommended for printing with PLA, it reduced the time required for the material to cool to its solidiﬁcation temperature [21].

Larger nozzle diameters increase the the length of time the material takes to cool, a 0.8 mm diameter nozzle was found to result in an acceptable cooling time, whilst being suﬃciently large to give printed parts stiﬀness.The use of layer-by-layer AM with diﬀerent orientations to the gravity vector has been investigated to determine the feasibility of using this technique in microgravity [22, 23]. It was found that a similar part quality can be achieved to conventional printing; however, to the author’s knowledge, support-free printing in these environments has not yet been explored. Therefore, to determine the eﬀects of modifying the print orientation to the gravity vector on support-free printing, a separate setup was used. This consisted of a ﬁxed six-axis UR3 robotic arm6, which has the same kinematics as the Meca500.

This consisted of a ﬁxed six-axis UR3 robotic arm6, which has the same kinematics as the Meca500. The same 3D printing equipment and printing parameters (0.8 mm diameter nozzle, print temperature of 200 °C) as used in the free-ﬂyer experiments were applied.2.3. Control SystemTwo GNC algorithms are used for controlling the movement of the free-ﬂyer on the air bearing table. The ﬁrst is a proportional–integral–derivative (PID) control algorithm, which is used for station keeping, i.e. when the free-ﬂyer should maintain a given position and orientation, whilst the robotic arm is moving. Three separate controllers regulate the 𝑥-, 𝑦- and 𝜃-position; however, they do not allow for velocity and acceleration control.

Three separate controllers regulate the 𝑥-, 𝑦- and 𝜃-position; however, they do not allow for velocity and acceleration control. A novel sliding mode controller (SMC) is used for navigation along trajectories, allowing simultaneous position, velocity and acceleration control [24].The free-ﬂyer is exposed to diﬀerent external disturbances that are caused by the system setup and are common to active air-bearing tables [25]. The most prominent is the residual acceleration due to an inclined table surface. The air cushion allows a frictionless planar movement which however, also leads to a movement of the free-ﬂyer if the table is not aligned perpendicular to the gravitational pull vector. Aerodynamic forces and torques induced by pressure variations of the active air blowing system play a subsidiary role. Measurements on ELISSA show that aerodynamic disturbances are negligible while residual accelerations can reach up to 9.6 mm/s2.

Measurements on ELISSA show that aerodynamic disturbances are negligible while residual accelerations can reach up to 9.6 mm/s2. Imperfections of the individual table panels as well as the transition between them may increase the friction coeﬃcient between the free-ﬂyer and the table surface. The inclination of the table panels changes slowly with a maximal rate of 0.05 °/m along a panel which can be counteracted by the free-ﬂyer control system. Table imperfections however, can cause a sudden change of frictional force between free-ﬂyer and table. The free-ﬂyer control system is thus exposed to external disturbances of diﬀerent magnitude and rates of change.The SMC is used for the continuous printing demonstration as shown Figure 2a in and therefore performs a constant velocity trajectory along the 𝑥-axis. The controller gains are tuned experimentally.

The free-ﬂyer exhibits a maximum deviation from the linear trajectory of 2 mm as visible in Figure 3a. The velocity components 𝑣𝑥 and 𝑣 𝑦 have maximum deviations of 1 mm/s and 1.5 mm/s from the target velocity proﬁle respectively during the acceleration phase. The deviations during the constant velocity phase are less than 0.5 mm/s (see Figure 3b).The PID control system is used for the segmented printing method (see Figure 2b). For each segment, the free-ﬂyer is held stationary while the robotic arm performs the printing movements. To achieve accurate results the free-ﬂyer has to counteract the aforementioned external disturbances, but also the induced reaction forces due to the robotic arm movement, as well as adhesive forces of the extruded ﬁlament. The PID controller gains are tuned experimentally.

The PID controller gains are tuned experimentally. A test of the PID controller without moving the robotic arm or extruding ﬁlament onto a substrate shows target position deviations of the free-ﬂyer of less than ±1 mm at maximum and a standard deviation of 0.4 mm.15153. Results and Discussion3.1. Variation of Print Orientation with Gravity VectorTo establish the feasibility of using the support-free printing method in microgravity, the eﬀects of varying the print orientation with respect to the gravity vector were investigated. Experiments were carried out using a robotic arm mounted to a stationary platform, which printed a support-free truss onto the substrate. This was performed with the substrate perpendicular to the gravity vector, once with the nozzle pointing downwards, i.e. in 1 g, and once with the nozzle pointing upwards, i.e. in −1 g. Figures 4a and 4b show the trusses printed in the conventional, and anti-gravity orientations respectively.

in 1 g, and once with the nozzle pointing upwards, i.e. in −1 g. Figures 4a and 4b show the trusses printed in the conventional, and anti-gravity orientations respectively. It can be seen that there is no visible diﬀerence between the two, conﬁrming that support-free printing is possible independent of the gravity direction, and thereby implying that it would also be suitable for a microgravity environment. Though no diﬀerence between the print performance was observed, further investigations will evaluate mechanical properties between trusses printed in the two orientations, as it is beyond the scope of this initial feasibility analysis.(a)(b)Fig.

4 A visual comparison of trusses printed with the print platform normal to the gravity vector, in (a) 1 and (b) -1 g, with the printing resolution and the nodal behaviour being comparable in both cases15162-10123 Target0100200300400500600700800Relative x-Position [mm](a)020406080100120140160180200220Time [s]202468Velocity [mm/s] Target vy(b)(c)Fig. 5 Comparison of (a) deviation from the target 𝑦-position, (b) deviation from the target velocities, and (c) extruded material when printing using the SMC controller to follow a constant velocity trajectory (beading highlighted by ellipses)15173.2. Continuous PrintNext, the capability of printing a structure by extruding whilst moving the robotic arm and the free-ﬂyer using SMC, as explained in Section 2.2, was investigated.

Continuous PrintNext, the capability of printing a structure by extruding whilst moving the robotic arm and the free-ﬂyer using SMC, as explained in Section 2.2, was investigated. The extrusion of a continuous truss element as displayed in Figure 2a requires the simultaneous control of free-ﬂyer position, robotic arm movement and ﬁlament extruder velocity in a closed-loop system. To evaluate the inﬂuence of further external disturbances due to the material extrusion, a continuous line was extruded using the same trajectory as in Figure 3, whilst keeping the robotic arm in a ﬁxed position.Figures 5a and 5b show the free-ﬂyer’s trajectory proﬁle in comparison to the targeted proﬁle. Deviations of position and velocity are of a similar magnitude compared to the initial test without extrusion (Figure 3). The results show that the extrusion of material has no further eﬀect on the free-ﬂyer. Figure 5c displays two extruded lines following the trajectory proﬁle of Figures 5a and 5b.

The results show that the extrusion of material has no further eﬀect on the free-ﬂyer. Figure 5c displays two extruded lines following the trajectory proﬁle of Figures 5a and 5b. The extrusion appearance is consistent indicating that small deviations of free-ﬂyer and extruder velocity are tolerable. Beading is only visible during the acceleration and deceleration phase by nature of the uncoupled free-ﬂyer and extruder velocities (highlighted in Figure 5c). The similarity of both extrusions demonstrate the repeatability of the SMC. The development of a closed-loop system incorporating the free-ﬂyer, robotic arm and FFF printer setup would allow the manufacture of a continuous truss as shown in Figure 2a and will be investigated in the future.3.3. Segmented PrintAs previously discussed, an alternative method for printing a large structure is to split it into segments, each of which are printed in turn with the spacecraft in a diﬀerent position.

In the case of a structure which follows a consistent pattern, such as a truss, it is possible to design a single segment which can be repeated until the desired size is achieved. The beneﬁt of using this method is the simpliﬁcation of the spacecraft’s control, where only the position needs to maintained, and not the velocity. As a spacecraft can move in 6 DOFs, these joints could be in any orientation, and would allow the manufacture of structures with unlimited dimensions in any direction.To investigate the feasibility of this manufacturing method, a design for a simple repeatable truss was created. As the ﬁrst segment was not required to join to an already present segment, it used a slightly modiﬁed design. Subsequent segments however, used an identical design. Figure 6a shows an example of the ﬁrst segment printed using a robotic arm mounted to a stationary surface.

Subsequent segments however, used an identical design. Figure 6a shows an example of the ﬁrst segment printed using a robotic arm mounted to a stationary surface. The excess material to the right of the truss is used to prime the nozzle, whilst the material on the left is where the next truss segment will be joined. As mentioned in the previous section, to allow support-free printing, the extrusion velocity must be lower than the printhead velocity; for these experiments, a printhead velocity of 3.33 mm/s was used, whilst an extrusion velocity of 2.75 mm/s was used. It should be noted that the design for each segment utilises a continuous print path, opening up the possibility of using continuous ﬁbre reinforcement in the future.Figure 6b shows the same design printed using a free-ﬂyer which was maintaining a ﬁxed position. It can be seen that though the basic shape remains the same, there is some inconsistency in the print (e.g. missing node attachments).

It can be seen that though the basic shape remains the same, there is some inconsistency in the print (e.g. missing node attachments). This is due to variations in the free-ﬂyer’s position during the print. As seen in Figure 7, the free-ﬂyer’s position varies by up to ±1 mm during a print, and it rotates up to ±1° (see Figure 7a). As the printer nozzle is located approximately 400 mm from1518(a)(b)Fig. 6 A comparison between printing a truss segment (a) using a robot arm mounted to a stationary surface and (b) using a robot arm mounted to a free-ﬂyer which is station keepingthe geometric centre of the free-ﬂyer, the angle error translates to a position error for the nozzle of ±6 mm in the 𝑥-direction, this is shown in green in Figure 7a. This angular error is partially due to the movement of the robotic arm. Though the arm moves slowly, and has a low mass compared to the total system, any movement by the arm induces a movement in the free-ﬂyer due to dynamic coupling.

Though the arm moves slowly, and has a low mass compared to the total system, any movement by the arm induces a movement in the free-ﬂyer due to dynamic coupling. Furthermore, during printing the print nozzle experiences friction with the print platform which, due to the lever arm, results in a moment being applied to the free-ﬂyer. To determine the extent to which the errors are due to dynamic coupling, and to what extent they are due to friction, the robotic arm was commanded to move through the same motions required to print a truss segment, whilst the extruder was turned oﬀ, and the print platform moved away. In this case the free-ﬂyer orientation deviated by ±0.2°, which is also the error seen when the robotic arm does not move.

In this case the free-ﬂyer orientation deviated by ±0.2°, which is also the error seen when the robotic arm does not move. It was therefore concluded that the error is due to friction experienced during printing, and that errors due to dynamic coupling are negligible.This rotational error is an important design consideration for a spacecraft operating in all 6 DOF as, depending on the mounting position of the robotic arm, all three axes of rotation may have a large impact on the position of the arm’s end eﬀector. It may be possible to reduce the impact of the error on the manufactured structure by having the robot arm actively compensate for any position error. However, a more robust control system coupled with reaction wheels, which are capable of1519generating a torque without residual linear forces, could eliminate this need.

However, a more robust control system coupled with reaction wheels, which are capable of1519generating a torque without residual linear forces, could eliminate this need. As the positional error introduced by the rotation of the free-ﬂyer linearly scales with the length of the lever arm, it is also advisable to reduce the lever arm length to a minimum.1-0.500.51Orientation Deviation [deg]0100200300400500600Time [s](a)8-6-4-202468 Free-Flyer Position Deviation4-2024 y [mm](b)Fig. 7 Deviation of free-ﬂyer whilst printing and station keeping, (a) shows the deviation in rotation, whilst (b) shows the deviation in 𝑥 and 𝑦 for both the free-ﬂyer and the nozzle15207654321100mmAfter the ﬁrst segment was printed, the free-ﬂyer was commanded to move to a second position 100 mm from the ﬁrst, before printing the second segment.

To ensure that the two segments successfully overlapped, and to compensate for height inconsistencies in the print platform, the origin of the print frame was meticulously set for the second segment. In the future, this could be achieved automatically using a camera and an object detection algorithm such as that used in [10]. Further segments were printed until a structure consisting of 7 segments was produced, as shown in Figure 8a, resulting in a total length of the truss of over 700 mm. By further repeating this process, it would be possible to print a truss with a length limited only by the size of the air bearing table. Figure 8b shows the joint between two truss segments, it can be seen that the material overlaps between the two segments.(a)(b)Fig.

Figure 8b shows the joint between two truss segments, it can be seen that the material overlaps between the two segments.(a)(b)Fig. 8 Long truss consisting of 7 joined segments, (a) shows the whole truss which is over 700 mm long, whilst (b) is a close-up view of the joint highlighted between sections 3 and 4It can be seen that the printing inconsistencies are larger for some segments than others. Segment 6 in particular experienced large inconsistencies. Some areas of the air bearing table experience larger external disturbances than others. Additionally, the compression of air by the active blower system generates a large amount of heat during operation, raising the room temperature over time, which impacts the cooling speed of the support free structures.

However, these inconsistencies are all attributed to this experimental setup and are not present in the space manufacturing environment.One of the advantages of our proposed manufacturing method is the capability to manufacture structures with unlimited dimensions in any direction. To illustrate this, an ’L’-shaped truss section was designed and printed, and an additional linear truss segment joined to it (see Figure 9). It can be seen that this could be further extended in both the 𝑥- and 𝑦-dimensions by adding more segments.1521Fig. 9 ’L’ shaped truss for 2 dimensional structures with extension truss joined on 12 cm markThe mechanical strength of these trusses is impacted by the strength of the joints at the nodes, and between segments. As shown in Figure 10a, these are currently formed by simply extruding a layer of material on top of already cooled material; the adhesion strength at these points is therefore low.

As shown in Figure 10a, these are currently formed by simply extruding a layer of material on top of already cooled material; the adhesion strength at these points is therefore low. Typically, inter-layer adhesion is improved by applying a force when extruding the second layer of material, this could be incorporated as shown in Figure 10b. However, as discussed, any friction experienced by the printhead will increase the position error of the free-ﬂyer, and jeopardise the consistency of the print. An alternative therefore, is to preheat the previously printed layer using a non-contact method, such as a laser (Figure 10c), which has been shown to improve inter-layer adhesion [26].Limitations of the setup notwithstanding, the capability to manufacture structures, in a friction-less environment, with dimensions larger than the workspace of the robotic arm has been demonstrated.

Joining repeatable sections together would allow limitless fabrication in any dimension, permitting the manufacture of large structures such as antennas and trusses on orbit. The use of support-free printing to extrude in 3D space in particular has been shown to be capable of printing trusses in 1 g, and −1 g environments, implying that large truss structures could be printed using this method in microgravity. Going forward, the orientation of the printhead during printing will be varied. This would allow more ﬂexibility in terms of the height of free-standing structures which could be printed, without the printhead contacting already printed sections. Additionally, the main advantage of 3D printing opposed to the manufacturing methods discussed in Section 1 is the possibility for limitless manufacturing in all three dimensions. Further experiments would therefore be extended to printing structures larger than the reach of the robot arm in two dimensions.

Further experiments would therefore be extended to printing structures larger than the reach of the robot arm in two dimensions. In the future the capability to scale this technique to produce larger structures will be tested. For example, the nozzle diameter1522(a)(b)(c)Fig. 10 Comparison of techniques for bonding material at nodes. (a) Shows the method currently used, where material is joined by extruding the top layer on top of the previous layer. (b) Shows how material can be joined by applying a force when extruding the top layer. (c) Shows how material can be joined by preheating the already extruded layercould be further increased if a method can be found to cool the ﬁlament suﬃciently quickly, this would allow the production of stiﬀer structures, and minimise the impact of sub-millimeter position errors.

Finally, the trusses printed did not have their mechanical strength tested; future experiments will investigate this, as well as the use of other materials, and methods such as continuous ﬁbre reinforcement to improve the mechanical properties of the printed structures. The impact of the joints between segments on the mechanical properties on the larger structure will also be investigated.4. ConclusionIn this paper, two methods for manufacturing large structures were explored using a robotic free-ﬂyer on an air bearing table: extruding whilst the free-ﬂyer was moving; and splitting the large structure into segments which could be printed in turn. It was found that the continuous printing method was capable of producing consistent results for a single line of material. Using the segmented printing approach, it was possible to print a truss structure with a length over 700 mm.

Using the segmented printing approach, it was possible to print a truss structure with a length over 700 mm. The dependency of the support-free printing method on the direction of the gravity vector was explored, and it was found through visual inspection that printing in −1 g produced no discernible diﬀerences with printing in 1 g. As such, it is posited that the method can be used independent of the gravity direction, and therefore also in microgravity. We have thus demonstrated the feasibility of a small spacecraft to manufacture support-free structures larger than itself using a robotic arm with 3D printhead end eﬀector. Using such a spacecraft would allow support-free 3D printing of structures with unlimited dimensions, and oﬀers more ﬂexibility than the truss manufacturing machines proposed by others.1523References[1] Gerard K. O’Neill. The colonisation of space. In Space Manufacturing Facilities, Reston,Virigina, 1975.

viewed PaperFEASIBILITY STUDY OF LARGE-SIZED ALUMINUM FACADES BY USING WIRE ARC ADDITIVE MANUFACTURINGY. Yamagata*, T.Sagawa*, M.Nitawaki*, and T.Abe †Institute of Technology, Shimizu Corporation, Japan † Graduate School of Science and Engineering, Saitama University, JapanAbstractOver the past few years, there has been growing interest in the fabrication of construction components by using wire arc additive manufacturing (WAAM). We focused on the finishing materials as a potential application for WAAM and began to consider fabricating aluminum building facades. However, there are several issues, such as fabrication size of 4-5m, aesthetics, and structural performance. Therefore, a trial fabrication and a non- destructive static loading test were conducted. In the study, an aluminum chair was used as a model because it contains the engineering basis of the building facades.

In the study, an aluminum chair was used as a model because it contains the engineering basis of the building facades. In the trial fabrication, a method of leveling the build surface for each of the multiple layers was found to be effective for large-sized fabrication. Bead blasting was also effective in removing oxides from aluminum surfaces and adjusting the appearance. In the loading test, both displacement and strain measurements agreed well with the FEM analytical values. The results showed that WAAM has the possibility of fabricating large-sized aluminum building facades with the structural performance expected in the FEM analysis.1. IntroductionThe computation-based approach in design, called computational design, has been developed over the last decade [1] and is becoming more common among architects and designers. This method is used to derive the optimum shape through repeated trials in the program.

This method is used to derive the optimum shape through repeated trials in the program. However, the optimum shape may not be bound by existing concepts and be too complex to be realized by conventional manufacturing processes, such as machining or casting. For this reason, additive manufacturing (AM), which offers a high degree of freedom in modeling, has been attracting attention in the construction industry in recent years [2]. AM is defined as “process of joining materials to make parts from 3D model data, usually layer upon layer, as opposed to subtractive manufacturing and formative manufacturing methodologies” in the international standard ISO/ASTM52900:2021 [3]. AM can produce complex or customized parts directly from digital models.

AM can produce complex or customized parts directly from digital models. Therefore, products can be manufactured in one step as designed without the constraints of conventional manufacturing processes.Regarding metal additive manufacturing, powder bed fusion (PBF) has been widely used for relatively high accuracy and high unit cost parts used in the automotive, aerospace, and medical industry [4]. In contrast, PBF has a limited range of applications in the construction industry. That is because construction components require a large size and low unit cost, whereas PBF has a relatively small build volume and high material and equipment costs. Thus, the use of directed energy deposition (DED) is being considered [5], which has a larger build volume and lower cost than PBF [6]. Especially, wire arc additive manufacturing (WAAM) is attracting attention among DED methods. WAAM is a method to form a target shape by layering weld beads using arc welding technology.

Especially, wire arc additive manufacturing (WAAM) is attracting attention among DED methods. WAAM is a method to form a target shape by layering weld beads using arc welding technology. Although WAAM is less accurate than other DED methods, it is superior in terms of build volume, cost, or mechanical properties [7], which has a high affinity with the construction industry. Recently, studies on the use of WAAM in the construction industry have been mainly focused on the fabrication of structural materials using steel [8][9]. However, there have been few studies on finishing materials using aluminum alloys, which are frequently used for building facades. Therefore, the purpose of this study was to investigate the feasibility of fabricating large-sized aluminum building facades by using WAAM.There were several issues in using WAAM for fabricating building facades.

Since aluminum alloys have a lower melting point and higher thermal conductivity than steel, the accuracy of the bead shape is prone to be lower due to temperature change during the deposition. This makes layer height variation from place to place and1465makes it difficult to continue deposition when the height in the building direction becomes high. Furthermore, aluminum alloys are more prone to occur oxidation than steel [10], and the components fabricated by WAAM leave an uneven surface. This may negatively affect their aesthetics and structural performance. Therefore, these issues were studied by using a chair as a model, which requires both an aesthetically pleasing appearance and structural performance that can withstand realistic use. For the study, a trial fabrication and a non-destructive static loading test were conducted.

For the study, a trial fabrication and a non-destructive static loading test were conducted. In the trial fabrication, a method of leveling the build surface for each of the multiple layers was adopted and evaluated to increase the stackable height. The appearance of the chair was also checked immediately after deposition and after surface treatment by bead blasting to confirm the aesthetics. In the loading test, measured displacement and strain were compared with the FEM analysis results to evaluate the structural performance.2. Design and FEM analysisA chair was designed as a simple model that includes the engineering basis of building facades. The chair size was 420mm (width)×400mm (depth)×600mm (height) with a hollow cross-sectional structure, as shown in Figure 1. The uniform cross-sectional plate (5.0 mm width) and joints (8.0 mm width) were designed, and the joints structurally united the two plates on the front and back sides.

The uniform cross-sectional plate (5.0 mm width) and joints (8.0 mm width) were designed, and the joints structurally united the two plates on the front and back sides. The chair material was Al-5356, which is widely used for welding.FEM analysis was performed for the preliminary evaluation of the structural performance. Table 1 shows the material properties of Al-5356 used in the analysis. The values were the nominal value of density [11] and the results of tensile tests on Al-5356 specimens fabricated by using WAAM [12][13]. The analysis was performed under the condition that a static load of 981N was applied to the seat surface. Figure 2 shows the FEM analysis results. The maximum total displacement was 29.9 mm near the top of the chair, as shown in Figure 2 (a). The maximum Mises stress was 80.4 MPa, and the stress was concentrated at the joints in the chair leg, as shown in Figure 2(b).

The maximum Mises stress was 80.4 MPa, and the stress was concentrated at the joints in the chair leg, as shown in Figure 2(b). The safety factor based on yield strength was 1.43, confirming that the chair was in the elastic range when a static load of 981N was applied.Table 1 Material properties of Al-5356 used in the FEM analysisParameters Density Young’s modulus Poisson's ratio Yield strengthUnit kg/m3 GPa - MPaValue 2640 68.4 0.335 115Reference [11] [12] [12] [13]A600 mmA420 mm981N[mm]0.018.018.029.9980N12.024.012.06.029.90.024.06.054.036.00.072.018.080.4981N[MPa](a) Total displacement(b) Von Mises stressFig.1 Chair designFig.2 FEM analysis results14663. Trial Fabrication3.1 Fabrication procedureTo increase the stackable height, the method of leveling the build surface for each of the multiple layers was adopted.

Trial Fabrication3.1 Fabrication procedureTo increase the stackable height, the method of leveling the build surface for each of the multiple layers was adopted. The leveling process was presumed to allow the deposition process to continue because the unevenness of the build surface would be removed. Figure 3 shows the schematic process sequence for the chair fabrication. Deposition of six layers, leveling the build surface with a grinding tool, measurement of the topmost bead width with calipers, and waiting for cooling were repeated. The deposition process was performed by robot- assisted welding. A Fronius CMT welding power source (TPS400i) combined with a Yaskawa Electric 6-DOF robot arm (MOTOMAN-AR1440) was used, as shown in Figure 4. Figure 5 shows the torch path for deposition. Since the deposited shapes were the same in all layers, the torch direction was always oriented vertically.

Figure 5 shows the torch path for deposition. Since the deposited shapes were the same in all layers, the torch direction was always oriented vertically. The x and y coordinates of the starting point of the torch path were the same, and the torch travel direction alternated between counterclockwise and clockwise for each layer. Table 2 shows the welding parameters for deposition. A substrate made of A5052 with dimensions of 1000 mm × 650 mm × 5 mm was clamped on a lattice table, and the welding beads were deposited on it. The leveling process and measurement process were performed by human hands.

The leveling process and measurement process were performed by human hands. The cooling process was performed by leaving the product at room temperature.Level the buildsurface with a grinding toolWait for coolingOne hour cooling timefrom previous depositionYesNoDeposit 6 layersMeasure topmost bead width with caliperTorchSubstrateTableChairCMTPower SourceShielding GasWire Feeder6-DOF Robot ArmFig.3 Process sequence for fabricationFig.4 Setup for robot-assisted wire arc additive manufacturing1cycleCWCWCCWStart pointCCWCWCCWyxyzFig.5 Torch path for deposition1467Table 2 Welding parameters for depositionParameters Wire material Wire diameter Shielding gas Shielding gas flow rate Current Wire feed speed Voltage Travel speedUnit - mm - L/min A m/min V mm/minValue A5356 1.2 Ar 18 120 8.1 14.8 18003.2 Result and discussionTo reach the stacking height of 420mm, 486 layers were deposited on the substrate.

The height of the deposited beads tended to be higher than the surrounding area at the joints. This was because the two deposited beads slightly overlapped due to the short distance of the torch path at the joints. Therefore, the leveling process worked effectively to remove the unevenness of the build surface. In this trial, the fabrication process was stopped when the stacking height reached 420 mm to match the chair design. However, it is possible that a stacking height of more than 420 mm could have been achieved if the fabrication process had been continued.Most of the beads at the joints were well joined to each other. However, the effective cross-sectional area of the joints was insufficient in rare cases, as shown in Figure 6.

However, the effective cross-sectional area of the joints was insufficient in rare cases, as shown in Figure 6. Figure 7 shows the topmost bead widths measuredNotjoinedJoint widthRare case (insufficient)Normal case (well joined)Deposited beadleveledFig.6 Joint conditionDesign width (plate)Measurement positionDesign width (joint)Layers [-]Topmost bead width [mm]Topmost bead width [mm]01002003004005006789101112010020030040050045678910Layers [-]Fig.7 Topmost bead widths measured for each of the six layers deposited1468Bead blastedBuilddirectionLarge unevennessFront sideBuilddirectionLarge uneven surfaceFront side of the chair leg(a) Immediately after deposition(b) After surface treatment (bead blasting)Fig. 8 Fabricated chair appearancefor each of every six layers deposited. The bead widths were wider than the designed widths and varied widely in the early layers.

8 Fabricated chair appearancefor each of every six layers deposited. The bead widths were wider than the designed widths and varied widely in the early layers. The average bead width at the plate was 6.7 mm (SD 0.87 mm) for layers 1-240 and 5.6 mm (SD 0.28 mm) for layers 241-486. The average bead width at the joint was 8.9 mm (SD 1.7 mm) for layers 1-240 and 8.3 mm (SD 0.34 mm) for layers 241-486. It can be presumed that the main reason for the increase and variation in bead width was the change in heat capacity during stacking. In the early layers, when the volume of the built object was small and the heat capacity was low, the temperature of the build surface may have been particularly easy to change. Consequently, the interpass temperature may have been high and varied widely, which caused the increase and variation in bead width.Figure 8 shows the fabricated chair appearance.

Consequently, the interpass temperature may have been high and varied widely, which caused the increase and variation in bead width.Figure 8 shows the fabricated chair appearance. The black soot-like deposits were formed on the surface of the chair immediately after deposition, as shown in Figure 8(a). The reason for this is presumed that the Mg elements contained in the wire were exposed to the outside of the shielding gas, oxidized, and adhered to the surface. To remove them, the surface of the chair was treated by bead blasting after the cut-off from the substrate. The oxides were removed and the metallic luster of the aluminum was brought out by this process, as shown in Figure 8(b). However, a large unevenness was observed on the surface of the early layers due to the trial and error of the workers who performed the leveling process.

However, a large unevenness was observed on the surface of the early layers due to the trial and error of the workers who performed the leveling process. This unevenness was aesthetically undesirable and needed to be resolved.The above results showed that the method of leveling the build surface for each of the multiple layers was effective for large-sized fabrication. However, since the deposited bead widths varied widely, control of interpass temperature should be studied further to deposit beads at the designed widths. Furthermore, bead blasting was also effective in removing oxides from the aluminum surface and adjusting its appearance. The aesthetics and shape accuracy could be further improved if the manual leveling process were automated by a robot.4. Non-destructive static loading test4.1 Test procedureTo evaluate the structural performance, displacement and surface strain were measured while static loads were applied to the seat surface of the chair.

A motion capture system was used to measure displacement, which can measure the three-dimensional movement of markers attached to the structure with an accuracy of better than 1 mm. Two digital image correlation (DIC) systems were used to measure strain, which are capable of non-contact surface measurement. Figure 9 shows the setup for the test. Seven cameras for the motion capture system and two sets of cameras for the DIC systems were installed around the chair. Spherical reflective markers were attached to the chair at a total of 12 locations, as shown in Figure 10(a), to measure the three-dimensional displacement of each marker. The measurement surfaces of the DIC systems were the front and back sides of the chair leg, as shown in Figure 10(b). Static loads were applied by placing plate-shaped weights (10kg×10) on the chair seat, as shown in Figure 10(c).

Static loads were applied by placing plate-shaped weights (10kg×10) on the chair seat, as shown in Figure 10(c). The load was placed in 98.1N increments by the weights, and when it reached 981N, it was unloaded in 98.1N decrements.1469Cameras for DIC 1Cameras for motion capture systemWeightsCameras for DIC 2ChairLights for DICFig.9 Setup for non-destructive static loading testyz:Marker123456789101112xxDIC 2(Back side)DIC 1(Front side)z200mmWeight(10kg×10)(a)Marker positions for motion capture (b) Measurement surfaces of DIC(c) Static load with weightsFig.10 Measurement and load application method4.2 Result and discussionMeasured values for displacement and surface strain were compared with the FEM analysis results. Figure 11 shows the displacement of Marker 1 for each load in the x- and z-directions. The displacements varied almost linearly with the loads, and the maximum absolute value of residual displacement after unloading was 0.2mm.

The displacements varied almost linearly with the loads, and the maximum absolute value of residual displacement after unloading was 0.2mm. Therefore, there was little residual displacement, and the chair was in the elastic range when the load of 981 N196785589392589981196Load [N]392FEMLoadUnload7850510152025020406080100Displacement [mm]98120-15-10-50020406080100Displacement [mm] FEMLoadUnload(a) x-direction(b) z-directionFig.11 Displacement of Marker1 for each load147005101520250123456789101112Marker Number [-]Displacement [mm]FEMMeasured20-15-10-5050123456789101112Displacement [mm] FEMMeasuredMarker Number [-](b) z-direction (a) x-direction Fig.12 Displacement of each marker when a load of 981N was appliedwas applied, as described in Chapter 2. The absolute values of displacement in the z-direction tended to be smaller than the analytical values.

The absolute values of displacement in the z-direction tended to be smaller than the analytical values. This is presumably because the bead widths were wider than designed, as described in section 3.2, and thus the stiffness increased.Figure 12 shows the displacement for each marker when a load of 981N was applied. The maximum displacement in the y-direction was 0.9 mm, which was sufficiently smaller than that in the x and z-directions, so the displacement diagram in the y-direction is omitted. In the x-direction, the maximum absolute error between the analytical and measured values was 2.88 mm at marker 3, with a relative error of 26.5%. In the z-direction, the maximum absolute error between the analytical and measured values was 0.98 mm at marker 2, with a relative error of 5.0%. In the error calculations, the analytical values were taken as true values.

In the error calculations, the analytical values were taken as true values. It can be presumed that the errors between the analytical and measured values in the large displacement region, where marker 1-6 were attached, were influenced by slight differences in force conditions. The displacements in the z-direction agreed well with the analysis because the fixation conditions of the chair were generally equal. The displacements in the x-direction deviated from the analysis compared to that in the z-direction because there were compound effects, such as the mismatch of fixing conditions of the ground contact, and the shift of the center of gravity of the load due to the inclination of the seat surface.Figure 13 shows the contour plots of analytical and measured values of strain in the vertical direction. The minimum and maximum values of the contour plots were taken as the minimum and maximum values of the analyzed strains.

The minimum and maximum values of the contour plots were taken as the minimum and maximum values of the analyzed strains. In the DIC1 measurement result, large strains were concentrated near the top of the measurement area. This seems to be an incorrect measurement because this area was the boundary between the chair leg and seat, where the distance between the DIC cameras and the measured object changed abruptly. The strains were consistent with the trend of the analysis, with large absolute values at the joints. However, the absolute values of the strains were larger than the analysis partially, such as the circled area in Figure 13. It can be presumed that the insufficient cross-sectional area of the joints, mentioned in section 3.2, may have caused stress concentrations. Since only the measurement of the uppermost width was performed in the trial fabrication, the distribution of joints with insufficient cross-sectional area could not be confirmed.

Since only the measurement of the uppermost width was performed in the trial fabrication, the distribution of joints with insufficient cross-sectional area could not be confirmed. However, if the internal shape is revealed, the causal relationship between shape and strain will become clear.We conclude that both displacement and strain measurements agreed well with the analytical values. Therefore, WAAM has the possibility of fabricating aluminum components with structural performance similar to that analyzed by FEM. However, since errors occurred between the analytical and measured values, it is necessary to set a safety factor to allow for this in practical use.

However, since errors occurred between the analytical and measured values, it is necessary to set a safety factor to allow for this in practical use. For a more detailed study, we should establish an inspection method to obtain internal shape, which is difficult to measure after fabrication.1471251100-1150[μm/m] Large strain area251100-1150 [μm/m]FEMDIC(a) Surface measured by DIC1 (Front side)DICFEM251100-1150[μm/m] Large strain area251100-1150 [μm/m](b) Surface measured by DIC2 (Back side) Fig.13 Contour plots of strain in the vertical direction5. ConclusionThere are several issues in using WAAM for fabricating building facades, such as fabrication size of 4- 5m, aesthetics, and structural performance. Therefore, the trial fabrication and the non-destructive static loading test were conducted to study these issues. In the study, the aluminum chair was used as a model because it contains the engineering basis of the building facades.

In the study, the aluminum chair was used as a model because it contains the engineering basis of the building facades. In the trial fabrication, the method of leveling the build surface for each of the multiple layers was adopted and evaluated to investigate the possibility of large-sized fabrication. The appearance of the chair was also checked immediately after deposition and after surface treatment by bead blasting to confirm the aesthetics. In the loading test, measured displacement and strain were compared with the results of the FEM analysis to evaluate the structural performance.

In the loading test, measured displacement and strain were compared with the results of the FEM analysis to evaluate the structural performance. The results showed the followings:(1) The method of leveling the build surface for each of the multiple layers is effective for large fabrication (2) Bead blasting is effective in removing oxides from aluminum surfaces and adjusting their appearance (3) WAAM has the possibility of fabricating aluminum components with structural performance similar to thatanalyzed by FEMTherefore, we conclude that WAAM has the feasibility of fabricating large aluminum building facades with the structural performance expected in FEM analysis. In other words, there is a possibility to fabricate large- sized and complex-shaped building facades at a relatively low cost.

In other words, there is a possibility to fabricate large- sized and complex-shaped building facades at a relatively low cost. Toward practical application in the construction industry, more research efforts should be conducted in the following 4 directions:1472(1) Increase in fabrication size aiming for 4~5m, which is the typical height of the floor (2) Control of interpass temperatures to improve the deposited shape accuracy (3) Installation of a robot for leveling process to improve the aesthetics (4) Establishment of an internal inspection method for more detailed analysis6.

asibility Study of small scale production based on additive manufacturing technologies¹ ² 3 Prof. Marlon Wesley Machado Cunico, Ph.D. , marloncunico@yahoo.com.br ¹Prof. Jonas de Carvalho, Ph.D. ¹Universidade de São Paulo campus São Carlos ²Concep3d Pesquisas Científicas 3Pontificia Universidade Católica do ParanáAbstract: Along the last years, the complexity of products has been growing progressively, while the product development life- cycle tended to be reduced. In addition to that, additive manufacturing technologies increased their role in the product development process, resulting in reduction of errors and products release time. In spite of these benefits, the main application of these technologies is still focused on initial phases of projects and results in high costs of parts and low volumes.

In spite of these benefits, the main application of these technologies is still focused on initial phases of projects and results in high costs of parts and low volumes. On the other hand, although conventional produtivity processes results in low costs and high volumes, the investiment related to these processes are high and the implementation time are long. For that reason, the main goal of this work is to investigate the possibility of application of additive manufacturing technologies for small and medium scale production. Along this work, the main direct and indirect processes which are used for small and medium scale production were studied and a numerical cost model were developed for each one. In order to compare the benefits and disvantages among the processes, 3 parts were selected and analysed through such models.

In order to compare the benefits and disvantages among the processes, 3 parts were selected and analysed through such models. By the end, the main cost, payback; amortization and takt time were identified and the most suitable process was found in accordance with annual part demand.Key words: Additive manufacturing, medium scale production, network production, flexible manufacturing systems1. INTRODUCTIONOver the last several years, the application of additive manufacturing (AM) processes has been steadily growing up as consequence of the advantages provided by it sort of process. In contrast with the benefits of these technologies, the main application is still focused on prototypes and special parts, as such medical devices (GIBSON et al., 2002; GIBSON et al., 2010; CUNICO e CARVALHO, 2013b; a). In parallel to those facts, the current production strategies are based on rapid or definitive tooling, resulting in high capital investments.

In parallel to those facts, the current production strategies are based on rapid or definitive tooling, resulting in high capital investments. As consequence, small scale production investments tend not to be justified or the payback time is too long (RUFFO et al., 2006).For that reason, the main goal of this work is to present a proposal of small scale production which is based on additive manufacture technologies.

As result, it is expected that the analysis and comparison of the current manufacturing process (injection moulding) versus 3 other additive manufacturing options indicates the solution that is more suitable for small scale production.In order to analyse the feasibility of additive manufacturing technologies as an effective production way, we established the injection moulding as the reference process in addition to studying 4 scenarios where the annual part demand, time demand, parts size and investment were the variables and the lead time, part cost, investments cost and pay-back period were the responses.In all the studied scenarios, we defined and indicated numerical models for the part cost estimation, where the definition of the main components of cost, lead time and minimal stock help to identify the feasibility of each scenario according to the part demand.In the first scenario, it was analysed the production feasibility of an injection moulding process where the injector machine and tooling costs were considering as amortised capital investment.

In this scenario, besides the analysis of feasibility, we have also presented an estimation model for the part, tooling and overhead costs, being useful for the process selection and the part cost estimation at the beginning of projects.In the second scenario, we investigated the feasibility of additive manufacturing services for production, where the costs related to production overhead and tooling are ignored. In this scenario, it is also important to see that besides the lead time and inventory dimensioning play a fundamental role in this business segment, these parameters might determine the feasibility of a new product release.In the third and fourth scenarios, we analysed two different production strategies where additive manufacturing technologies are considered. In both cases, the acquisition of equipment was considered in the investment and part cost estimation.

In both cases, the acquisition of equipment was considered in the investment and part cost estimation. At the end, all the scenarios were compared in order to identify suitability the production strategy for small and medium scale strategies.In addition to the feasibility analysis of small scale production products, this work can also be a very useful tool for customised or tailor-made business segments, where the feasibility of new products is hardly achieved and the product cost tend to be extremely high.2.

PRODUCTION ESTIMATION MODELSIn order to investigate and compare the part cost of parts which are made in conventional injected mould and additive manufacturing techniques, we selected the main components of the part cost and created estimation cost models of these components.In general way, the main part cost components can be classified in direct and indirect costs, as shown in Figure 1 (REES, 1996; CHATAIN e DOBRACZYNSKI, 1997; ASIEDU e GU, 1998; FAGADE e KAZMER, 2000; ROSATO e ROSATO, 2000; NIAZI et al., 2006; RUFFO et al., 2006).

Nevertheless, as the main goal of this work is to compare different production scenarios, we excluded the administrative overhead from the analysis.In this way, it is possible to see that the direct cost is related to the material which is directed used to fabricate the part, while the indirect costs concern the process time, labour, investments and are amortised by the volume of fabricated parts.Figure 1 – Representation of the main part cost components(For this study, the main cost components that have influence in the part cost were analysed, as such direct cost ). Therefore, thetdirectCcos_toolingCmachineCoverheadC), tooling cost (), machine cost () and production overhead cost (part cost (partC) might be defined as:machineoverheadtoolingdirectpartCCCCCIn addition to the analysis of part cost, it was also analysed the feasibility of the production scenarios with respect to demand and lead time.

In fact, this is an important parameter to be analysed because it indicates whether the productive way is feasible, in addition to indicating the minimal stock which is necessary for each annual demand and the part demand time.For the definition of part demand time, we assumed that the annual demand is distributed homogeneously along theyear. Therefore, it was possible to see that the part demand time might be characterised by:annualjourneydailydaysworkingdemandNNNt__60In general way, the stock flow might be analysed by the variation of delivery parts (deliveryN) per lead time (and demand parts (demandN) per demand time (demandt), as shown in Eq.

(3)(1)(2)leadt)demanddowndemandleaddowndeliveryttroundNttroundNStock(3)Therefore, if we assume that the demand time is equal to the delivery time per part, it is possible to estimate the minimal inventory which is needed to attend production through the maximum of stock curve in addition to the safety stock.2.1.INJECTION MOULDINGIn order to identify the total cost of an injected part, we applied an estimation cost model to identify the tooling cost, while the machine cost was obtained by quotation.

The part direct cost was estimated through the part volume and the raw material cost was identified by low volume quotation.DIRECT COSTThe direct cost of part is related to the quantity of material which is necessary to fabricate the part, where the ) andpartVmaterialpart_materialpartk_volume of part (), the specific weight of material (), the raw material coefficient (the material waste (wasteC) define the direct cost part (directC).wastematerialpartmaterialpartpartdirectCkVC__(4)TOOLING COSTFor the estimation of tooling cost, we applied Boothroyd and Dewhurst’s method, which concern in the estimation of operational and direct costs which are necessary to build a cold runner mould (REES, 1996; CHATAIN e DOBRACZYNSKI, 1997; ASIEDU e GU, 1998; FAGADE e KAZMER, 2000; ROSATO e ROSATO, 2000; WANG et al., 2003; CAMPO, 2006; NIAZI et al., 2006; FONSECA et al., 2007; KAZMER, 2012).

In this model, the main inputs which are used to identify the total cost of injection mould are the volume of part and number of cavities.For this estimation, it is possible to separate the total cost of mould (toolmouldingC_) in three main components:Cavities cost (cavitiesC), mould base costs (basemouldC_) and customisation costs (ioncustomisatC), as presented in Eq. (5)(REES, 1996; CHATAIN e DOBRACZYNSKI, 1997; FAGADE e KAZMER, 2000).ioncustomisatbasemouldcavitiestoolmouldingCCCC__(5)For the specification of cavities cost, it is established the individual cost of each mould cavity (cavitiyC) multipliedby the number of cavities (cavitiesn) and a discount factor (discountcavityf_), as it is possible to see in Eq. (6). In thisstudy, we ignored the discount factor per cavity in addition to establishing that the number of cavities should be limited to 5.

(6). In thisstudy, we ignored the discount factor per cavity in addition to establishing that the number of cavities should be limited to 5. This restriction in the number of cavities was defined because the mould was design to low production volumes (REES, 1996; CHATAIN e DOBRACZYNSKI, 1997; FAGADE e KAZMER, 2000)..discountcavitycavitiescavitiycavitiesfnCC_(6)With respect to the cavity set cost, it is possible to identify that the main cost components are related to cavity ), as shown inmaterialcavitiyC_machiningcavitiyC_finishingcavitiyC_material (), cavity machining () and cavity finishing (Eq.

(14). In spite of the effect of thisfactors effect on the time estimation, we established that the efficiency of machining is high and the complexity of part was low, resulting in those factors being equal to 1 (REES, 1996; CHATAIN e DOBRACZYNSKI, 1997; FAGADE e KAZMER, 2000)..complexitycavitymachiningefficiencymachiningsurfacecavitiyvolumecavitiymachiningcavitiyfffttt_____(14)In addition, the estimation of the cavity volume machining time can be identified by the volumetric mould materialremoval rate, Eq.

(16).:speedfinishingmaterialcavityvolumematerialmaterialcavitiysurfacecavitiyFdARAt15.0____With respect to mould base cost and customization costs, we defined that these cost are respectively 15% and 150%of cavities cost.As the tooling cost is considered a capital investment, the contribution of tooling for the part cost is amortised by the volume of parts which is fabricated. Therefore, the total tooling cost might be defined as a function of moulding tool cost ( ) and annual parts amount (N), as shown in (17) (REES, 1996; CHATAINtoolmouldingC_df), depreciation factor (e DOBRACZYNSKI, 1997; FAGADE e KAZMER, 2000)..dtoolmouldingtoolingfNCC_PRODUCTION OVERHEAD COSTWith reference to the estimation of basic production overhead costs (Eq.

(19). In general, it was alsocyclemouldingt_setuptdirect moulding cycle time () and setup time (considered that the cycle time is around 30s, as the injected parts moulding cycle commonly varies from 15 to 120 seconds (REES, 1996; CHATAIN e DOBRACZYNSKI, 1997; FAGADE e KAZMER, 2000).setupcavitiesBatchcyclemouldingbatchtnNtt_In addition, the manufacturing rate is related to the cost rate of machine operation (machineR) and labour (as it is possible to be seen in:machinelabourrateingmanufacturRRR_MACHINE COST(16)(17)(18)(19)labourR),(20)With respect to the machine and equipment acquisition (machineC), it was considered that the total cost ofequipment (equipmentC) is amortised by the amount of parts which is fabricated during a payback period of time(paybackT). In this case, the total amount of parts is defined by the annual demand of parts (annualN) multiplied by thepayback period, as it is possible to be seen in Eq.

(36) (REES, 1996; CHATAIN e DOBRACZYNSKI, 1997; FAGADE e KAZMER, 2000; ROSATO e ROSATO, 2000; KAZMER, 2012).paybackannualequipmentmachineTNCC(21)In addition, one of the most commonly methods for equipment specification was the clamp tonnage method, whichidentify the machine clamp force through the number of cavities, projected area, mould size and shot capacity.For the determination of shot capacity (capacityShot), we can use the weight of part (partWt), number of cavities(cavitiesn) and sprue and runner factor (SRF) (ROSATO e ROSATO, 2000; KAZMER, 2012).)28.37(g/oz)/(4545.1)(16)()(lbgozSRFngWtgShotcavitiespartcapacity(22)For hot runner mould systems,SRFis equal to 1, while for cool runner systemsSRFis characterised by :15.15.0partWtSRF(23)It is also possible to identify the melt capacity of machine through:)(30)()(stgShotgMeltCyclecapacitycapacity(24)Through this parameter, it is possible to identify the clamp force of machine in a preliminary way according to theFigure 2.Figure 2 – Machine shot capacity versus machine clamp force (ROSATO e ROSATO, 2000)On the other hand, it is also possible to estimate the machine clamp force through the wall thickness method.

In this ) and a wall thicknessclampFprojectedAmethod, the clamp force () is found by the projected area of runner and cavities (factor, which can be seen in the Eq. (25).thicknesswallprojectedclampfAtF_)(Table 1 – Investigation matrix of additive manufacturing services wall thickness (in) Wall thickness factor (t/in²)0.020-0.062 0.062-0.125 0.125-0.2506-5 5-4 4-3As consequence, the general estimation of machine cost might be estimated by the update regression of(BOOTHROYD e DEWHURST, 1988):)(43016000tFCclampequipment2.2. ADDITIVE MANUFACTURING SERVICESFor the study of additive manufacturing services as productivity way, we have also considered that the tooling cost, machine cost and production overhead were null, while the direct cost remained the responsible component of part cost. It is important to note that there are other costs inherent to this kind of scenario, as such logistics, stock and quality.

It is important to note that there are other costs inherent to this kind of scenario, as such logistics, stock and quality. Nevertheless, these costs were ignored in this study in order to create a comparison criterion among the studied scenarios (RUFFO et al., 2006).In order to identify the cost estimation of additive manufacturing services, it was requested quotation of 3 dimension parts, 3 technologies, and 4 part quantities in order to be obtained a statistical regression and the cost estimation formulation.

The investigation matrix used to analysed the service cost estimation can be seen in Table 2, where it is related the part quantity per order, main dimensions of analysed parts and fabrication technology.(25)(26)Table 2 – Investigation matrix of additive manufacturing servicesIn addition to being identified the regression equation for each technology and the part cost estimation model, it was also determined the minimal stock volume which would be needed to attend annual part demand. Therefore, the feasibility of this productivity concept might be evaluated.It was also find that the lead time was a based on service time that the bureaux provide (batcht) and the shippingtime (shippingt), as presented in Eq.:(27). Therefore, the minimal inventory might be estimate in accordance with themaximum value of Eq. (3).shippingbatchleadttt(27)2.3.

Therefore, the minimal inventory might be estimate in accordance with themaximum value of Eq. (3).shippingbatchleadttt(27)2.3. ADDITIVE MANUFACTURING PRODUCTIONFor the generalised cost estimation of additive manufacturing parts, it was analysed the direct cost and production )overhead as a function of part size, building area, parts demand and batch volume. Therefore, the part cost (partCmight be characterised two conditions: single part batch and optimised batch. In spite of both conditions being defined by Eq.

Therefore, the part cost (partCmight be characterised two conditions: single part batch and optimised batch. In spite of both conditions being defined by Eq. (28), the overhead cost is amortised by the number of parts per batch, resulting in part cost differences (ASIEDU e GU, 1998; HOPKINSON e DICKENS, 2001; HOPKINSON e DICKNES, 2003; RUFFO et al., 2006).machineoverheaddirectpartCCCC(28)In this study, the estimation of direct cost was mainly determined by the cost of material, whose main components ) and waste ), material density (materialpartk_materialpart_partVare the raw material cost coefficient (), part volume (material (wasteC), as shown in Eq. (29).

(29). In addition, we have also considered that the waste material is 10% of partmaterial because of the support material, errors, purges routine, among others.wastematerialpartmaterialpartpartdirectCkVC__(29)In this case, it is important to note that the part volume consider a solid strategy, being ignored either weave infill,pattern infill, low density or airgap strategies (GIBSON et al., 2002; GIBSON et al., 2010).With reference to production overhead costs (Eq. (30)), we defined manufacturing batch time (batcht),manufacturing rate (rateingmanufacturR_) and the amount of parts per batch (batchN) as the main cost components. It isimportant to be highlighted that although additive manufacturing processes do not result in amortised tooling cost, the operational cost is amortised by the number of parts which is produced in each batch.

As consequence, the maximum amortisation is restricted by the building area of machine.batchrateingmanufacturbatchoverheadNRtC_(30)In this case, the determination of the maximum number of parts which are possible to be produced in each batch andpartLpartWareabuildingL_was established by the maximum dimensions of part (and), building area (areabuildingW_) and the minimal distance between parts (s), resulting in:sWWroundsLLroundNpartareabuildingdownpartareabuildingdownbatch__max_(31)It is also important to note that this equation ignores the possibility of building several parts along z axis, resultingin a bidirectional part building matrix.For the determination of total batch time, we defined that the main components that contribute for the batch are thedirect manufacturing time (ingmanufacturt), setup time (setupt) and post-processing time (gprocesposttsin), as represented:gprocespostsetupingmanufacturbatchleadtttttsin(32)We have also defined that the manufacturing time is a result of the building layer height (layerH), part height(partH), part length (partL), part width (partW), number of parts per batch (batchN), raster feed rate (rasterF), delaytime per change of layer (delaylayert_) and the tool diameter (d), as shown in Eq.

(33). In this case, the tool diametermight represent either nozzle diameter, bead width or laser beam diameter, while the part infill strategy was considered solid.delaylayerrasterbatchpartpartlayerpartingmanufacturtFNdWLHHt_(33)Additionally, we can also define the average lead time (partt) per part as:batchbatchpartNtt(34)On the other hand, the manufacturing rate (rateingmanufacturR_) was defined to be mainly compounded by theoperational time-machine cost rate (machineR), energy cost rate (energyR) and labour cost rate (labourR), as presented inEq. (35).labourenergymachinerateingmanufacturRRRR_(35)By the end, the last component of part cost is related to the machine and equipment acquisition (machineC). As thiscost is a capital investment, the total cost of equipment (equipmentC) is amortised by the amount of parts which isfabricated during a payback period of time (paybackT).

As thiscost is a capital investment, the total cost of equipment (equipmentC) is amortised by the amount of parts which isfabricated during a payback period of time (paybackT). In this case, the total amount of parts is defined by the annualdemand of parts (annualN) multiplied by the payback period, as it is possible to be seen in Eq. (36).paybackannualequipmentmachineTNCC(36)With respect to the equipment cost, we identified the approximated cost of the main professional additive manufacturing machines which presented building area superior to 300x300x300mm. In this case, we considered the FDM, SLA and SLS as the main technologies to be analysed in this scenario.2.4. LOW COST ADDITIVE MANUFACTURING IN NETWORKOn the other hand, the last scenario which we analysed in this work concerns the use of low cost additive manufacturing technologies as an effective way of production.

In contrast with the previous scenario, this proposal is marked by the use of low cost machines, which are also known as 3d printer.In general way, the main difference between the cost estimation model of this proposal and the additive manufacturing scenario is related to the possibility of simultaneous batches in addition to machine cost reduction. As consequence, it was possible to find that the production overhead and machine costs were the most affected component in the model.Adjusting the previous model to the number of machines (machinesN), it can be seen that the machine cost is:paybackannualmachinesmachineindividualmachineTNNCC_(37)For this scenario, we identified the approximated cost of the main low cost additive manufacturing machines whichpresented building area up to 150x150x150mm. In this case, we considered only the FDM, SLA technologies.(And the labour cost rate might be amortised by the number of machines which one worker can manage ). In Eq.

In this case, we considered only the FDM, SLA technologies.(And the labour cost rate might be amortised by the number of machines which one worker can manage ). In Eq. (38), it is possible to see the formulation of this cost rate as a function of number of machineslabourmachineN_used in network (machinesN), number of managed machines per worker (labourmachineN_) and cost of worker per hour(labourC).labourlabourmachinemachineslabourCNNR_(38)In addition, it was also possible to find the variation of batch lead time and the average part lead time for more than ), as presented in Eq. (39) andmachinesNone batch as a function of the number of machines which is used in network (Eq.: (40). On the other hand, the lead time for the first batch is equal to the batch time.machinesbatchleadNtt(39)machinesbatchmachinesbatchbatchpartNNNttt(40)3.

On the other hand, the lead time for the first batch is equal to the batch time.machinesbatchleadNtt(39)machinesbatchmachinesbatchbatchpartNNNttt(40)3. RESULTS AND DISCUSSIONS3.1.INJECTION MOULDINGWith respect to the results of this study, it was possible to characterise the cost of part as a function of annual part demand in addition to production strategy. In order to be possible to compare all the 4 studied scenarios, we established 3 main sizes of parts to be analysed: 8x8x15mm; 30x30x15mm and 60x60x15mm.For the cost estimation of moulded parts, we considered that the number of cavities should be equal to 4, which is an indicated value for low scale production.

The main parameters which were used in this analysis can be seen in Table 3, where the total cost of material, machine and tooling is also presented for each one of the 3 part sizes.Table 3 – Parameters for cost estimation of injection moulding partsIt is also important to note that the number of parts per batch directly interfere in the production overhead cost, as it is possible to see in Figure 3. In this figure, we can indicate the saturation of cost for batch sizes which are superior to 100 parts. For that reason, we selected the batch size equal to 100 parts to perform the cost analysis.Figure 3 – Effect of batch size for the production overhead costAs result of this cost analysis, it was possible to identify the variation of injection moulding part cost as a function of annual part demand, as shown in Figure 4. In this figure, we can also see the cost of the 3 part sizes in addition to the needed capital investment for each one.

In this figure, we can also see the cost of the 3 part sizes in addition to the needed capital investment for each one. In this case, besides the investment has varied from $16,000.00 to $65,000.00, the amortisation cost happens in exponential proportion. Therefore, if the $16.00 was considered a suitable part cost, the annual demand that justified the injection moulding production would be 1000, 3000, 5000 parts for each one of part sizes.Figure 4 – Injection moulding part cost versus annual part demandAs the injection cycle defines the main lead time of parts, we can see the variation of demand and lead time for the parts, as shown in Figure 6. In this figure, it is possible to see that all the 3 analysed parts resulted in the lead time lower than the demand time.

In this figure, it is possible to see that all the 3 analysed parts resulted in the lead time lower than the demand time. As consequence, it indicates that the machine tend to present idleness for annual part demand inferior to 200000 parts/year.Figure 5 – Delivery and demand time as a function of annual part demandIn addition, if only one idle injection moulding machine were considered, the minimal inventory for200000parts/year demand and no safety stock would be 26.3.2. ADDITIVE MANUFACTURING SERVICESIn the second analysed scenario, we investigated the feasibility of additive manufacturing services as production way.

ADDITIVE MANUFACTURING SERVICESIn the second analysed scenario, we investigated the feasibility of additive manufacturing services as production way. In this study, we identified the tendency of cost which is related to parts as a function of number of parts per order or batch, as it is possible to be seen in Figure 6.In this figure, the diagrams of part cost as a function of number of parts per order or batch were presented in addition to a general diagram which compile the maximum, minimum and mean values of all the three analysed technologies.It is possible to see that the cost tends to be saturated in 50 parts batch sizes, while the size of parts proportionally increases the part cost.

For the FDM technology, values varied between $75.00 and $25.00 for 50 parts batch size, while the value remained near to $200.00 for a single part batch size.On the other hand, SLA part cost varied from $125.00 to $275.00 for one part batch size and from $17.00 to $50.00 for 50 parts batch size. It is also possible to establish as a general rule that SLA parts with main dimensions from 8 to 60 and 15mm of height mm tend to cost $30.00.For the SLS technology it was observed similar cost behaviour, where the cost varied according to the part size from $75.00 to $200.00 for a single part batch size. While for 50 parts batch size, the variation of cost was found between $11.00 and $75.00.Figure 6 – Diagrams of part cost estimation as a function of the number of parts per order for SLA, FDM and SLS technologiesAnother important point that was also seen in this study is related to stock analysis.

It was observed that the delivery time for additive manufacturing services is around 7 working days. Therefore, the delivery time per part is approximately 68 hours. In addition, Figure 7 presents the correlation of demand and delivery part time as a function of annual part demand.In this figure, it is possible to see that use of additive manufacturing services is feasible to be applied for annual partdemand inferior to 1715 in terms of lead time.Figure 7 – Delivery and demand time as a function of annual part demandWith respect to stock analysis, the minimal stock with no safety stock should be equal to 50 parts, if an annual partdemand of 1715 and a batch size equal to 50 were considered.3.3. ADDITIVE MANUFACTURING PRODUCTIONFor the Additive manufacturing production scenario, we estimate the part cost for 3 part sizes and 3 technologies, as it is possible to be seen in Table 4.

ADDITIVE MANUFACTURING PRODUCTIONFor the Additive manufacturing production scenario, we estimate the part cost for 3 part sizes and 3 technologies, as it is possible to be seen in Table 4. It is important to note that the machine cost is a quotation average and reflect the magnitude cost of each technology.In this table it is also exposed the maximum number of parts that might be produced by batch according to abuilding area equal to 300x300x300mm.Table 4 –Parameters for Cost estimation of additive manufacturing partsAnother point that is also important to note is related to the batch size, as presented in Figure 8. It is possible to see the tendency of cost saturation for batch sizes superior to 5 parts.

It is possible to see the tendency of cost saturation for batch sizes superior to 5 parts. In this figure, the variation of overhead cost as a function of part size is also presented, where the variation of cost is found between $4.00 and $45.00.Figure 8 – Effect of batch size for the production overhead cost of additive manufacturing productionIn addition, as the final part cost of additive manufacturing production was influenced by the amortised value of machine, we identified the variation of this cost as a function of machine technology and annual part demand, as presented in Figure 9.

According this figure, it is possible to see that the part cost tendency is around $25.00 for annual part demand superior to 10000 parts with dimension equal to 30x30x15mm.It is also presented that for this same annual part demand, the part cost of 8x8x15mm parts tend to $4.00 and60x60x15mm parts tend to $48.00.Figure 9 – Diagrams of part cost estimation as a function of the annual parts demand for SLA, FDM and SLS technologiesWith respect to the timing analysis, we have also estimated the maximum manufacturing time per batch according to the part size. In contrast, Figure 10 presents the comparison between the lead time and demand time as a function of annual part demand. In this figure, it might be indicated the manufacturing feasibility with accordance with annual part demand.

In this figure, it might be indicated the manufacturing feasibility with accordance with annual part demand. In this way, the maximum annual part demand that can be provided by additive manufacturing production in the studied conditions might be 1255 parts with 60x60x15mm, 2668 parts with 30x30x15mm and 15000 parts with 8x8x15mm.Figure 10 –Demand time and part lead time as a function of the annual parts demandWith respect to stock analysis, the minimal stock with no safety stock should be equal to 16 parts for 60x60x15mm part sizes, 81 part for 30x30x15mm part size and 900 part for 8x8x15mm part size, as it is possible to be seen in Figure 11. In this analysis, the maximum annual part demand for each part size was considered according to the presented before.Figure 11 –Stock flow of Additive manufacturing production as a function of part size3.4.

LOW COST ADDITIVE MANUFACTURING IN NETWORKNow for the last scenario, we analysed the feasibility of production which used low cost additive manufacturing machine in a network arrangement. In this case, the main parameters which were used can be seen in Table 5, where the machine cost and the building area dimensions are the main difference from the previous scenario. It is important to note that in this table, the production overhead cost considers only one machine in the estimation. Otherwise, the production overhead cost tends to decrease in accordance with the number of machines in the network arrangement.It is also possible to see that in comparison with professional additive manufacturing equipment, low cost equipment implied on an extremely high production overhead cost for large parts.

It probably occurs because of the low raster speed and the long manufacturing time.Table 5 –Parameters for Cost estimation of low cost network additive manufacturingRegarding Figure 12, it is possible to see the part cost as a function of number of machine in network and annual part demand. This figure indicates the decrease of part cost according to the rise of machine number in production network. It is also possible to evidence that the application of network arrangement make possible to reduce the part cost of 60x60x15mm in almost 7 times.Otherwise, the increase of machines into the network arrangement is not indicated for very small demand.

At this way, fabrication of very small parts (8x8x15m) with a single machine tends to be more interesting for annual demand which is found below 1000 parts, while the small parts (60x60x15mm) seems to be more suitable to be fabricated in a network when an annual demand is higher than 200 parts/year.Figure 12 – Diagrams of part cost estimation as a function of the annual parts demand for the number of low cost machinesIn Figure 13, it is presented the correlation between demand time as a function of annual part demand and the lead time which is provided by 1, 4, 8 and 16 machines in the network arrangement in addition to exposing the effect of part size for the lead and demand time.

With these diagrams, it is possible to identify the production way feasibility range, where the intersection between demand and lead time marks the maximum annual demand that the production network can support.Figure 13 – Demand time and part lead time as a function of the annual parts demand and number of machinesOn the other hand, the analysis of part lead time has shown to be strongly influenced by the number of machines in the arrangement. In this case, the maximum annual demand of 60x60x15mm part size that might be attended by the production way varied from 450 to 6000 parts/year if the number of machines in the arrangement would be increased from 1 to 16 machines.

In addition, for 8x8x15mm and 30x30x15mm part size, this number would respectively be raise to 80000 and 15000parts/year.With respect to the minimal stock considering no safety stock, this scenario implied on an inventory size equal to 4parts for 60x60x15mm part size, 16 for 30x30x15mm part size and 255 for 8x8x15mm part size.3.5. PRODUCTION STRATEGY COMPARISONComparing the results of the analysed scenarios, it was possible to identify the main differences among the scenarios in term of cost. In this analysis, it was also possible to see which production way is more suitable for each annual part demand.In order to compare the four scenarios, we identified the part cost of each process as a function of annual demand and part size, as represented in Figure 14. In this figure, it is possible to see that the most indicated production way for 30x30x15mm parts size and annual demand inferior to 1000 parts/year might be additive manufacturing services.

In this figure, it is possible to see that the most indicated production way for 30x30x15mm parts size and annual demand inferior to 1000 parts/year might be additive manufacturing services. On the other hand, for annual demand between 1000 and 3000 parts/year, the recommended production way should be additive manufacturing in a 8 machine network arrangement. In this case, it was also evidenced that injection moulding was the most indicated for annual demand superior to 3000 parts/year.For 8x8x15mm part size, it was indicated that additive manufacturing services is the most indicated production wayuntil 2000 parts/year, in addition to the network arrangement was seen to be equivalent to injection moulding part cost.In contrast with this, it was found that additive manufacturing services was the most indicated for 60x60x15mm part size with annual demand inferior to 500parts/year.

While the low cost additive manufacturing in 8 machine network arrangement was evidenced to be the most suitable for annual demand between 500 and 3000parts/year.

For superior values of annual demand, the most indicated process was proved to be injection moulding production.Figure 14 –Part cost as a function of annual demand for injection moulding, additive manufacturing services, additive manufacturing production and low cost additive manufacturing in network arrangementIn addition, the part demand time as a function of annual demand was also compared in Figure 15, where all the production scenarios were shown to be feasible in term of lead time for 8x8x15mm part size and annual demand inferior to 5000part/year.Figure 15 –Part demand time as a function of annual demand for injection moulding, additive manufacturing services, additive manufacturing production and low cost additive manufacturing in network arrangementOtherwise, the lead time of 30x30x15mm part size of additive manufacturing services was indicated to attend to 1500parts/year, while additive manufacturing production was to 2500parts/year.

For this part size, both injection moulding and low cost additive manufacturing in 8 machines network arrangement were found to support to annual demand superior to 5000parts/year.By the end, as the lead time of 60x60x15mm part size tend to be longer than smaller parts, the additive manufacturing production was found to attend to 1250parts/year, while additive manufacturing services was to 1500parts/year. Additionally, low cost additive manufacturing with 8 machines in a network arrangement was identified to support to 3000parts/year. For this part size, the only process that was found to attend to the demand time for annual demand superior to 3000parts/year was the injection moulding.With respect to the minimal inventory, we can see in Figure 16 that injection moulding result in the smallest inventory for small parts, while the additive manufacturing with 8 machines in a network arrangement does for medium size parts.

Moreover, although injection moulding and additive manufacturing services were found to imply in a constant inventory size for different part sizes the average of inventory size was marked to remain below 50 parts. In other words, no significant benefits in using additive manufacturing were seen for low scale production in comparison with conventional processes.Figure 16 –Minimal inventory for the maximum productivity4.

CONCLUSIONSIn this work, it was possible to see the main differences among injection moulding, additive manufacturing services, additive manufacturing with large professional machines and additive manufacturing with low cost machines in a network arrangement.In addition, estimated cost numerical models of each one of the analysed processes were developed and identifiedthe main components that contribute for the part cost.It was possible to evidence the feasibility range of each one of the analysed processes as a function of annualdemand besides being indicated the most suitable processes in accordance with the demand range.For very small demand, the most indicated production way which was found is additive manufacturing services even though it results in a high part cost. In contrast, low cost AM machines in a network arrangement were shown to be the most recommended for annual demand between 500 and 3000parts/year.

In contrast, low cost AM machines in a network arrangement were shown to be the most recommended for annual demand between 500 and 3000parts/year. It might indicate that this range of demand which was poorly covered may be attended by this proposed production way so that new business can also be created as consequence of this.With respect to the lead time analyses, it was evidenced that injection moulding attend all the analysed part sizes for annual demand superior to 5000parts/year, while the additive manufacturing services was up to 1500parts/year. On the other hand, additive manufacturing with 8 machines in a network arrangement was found to attend to annual demand superior to 5000parts/year for 8x8x15mm and 30x30x15mm part size and annual demand is limited to 3000parts/year for 60x60x15mm part size.5.

asibility study on plastic laser sintering without powder bed preheatingToshiki NIINO, Hisashi HARAGUCHI, Yutaro ITAGAKI,Institute of Industrial Science, the University of Tokyo4-6-1 Komaba Meguro Tokyo, 153-8505 Japanniino@iis.u-tokyo.ac.jpShintaro IGUCHI and Masashi HAGIWARAASPECT Inc.ABSTRACTIn a plastic laser sintering machine, the most part of power consumption is spent onheating of powder bed. The powder bed heating is essential to prevent parts fromwarping during the process. However metal laser melting is normally performedwithout such heating. During the process, warping is suppressed by fixing the parts to abase plate. In the present research, the same scheme was introduced to plastic lasersintering. A plate of 2mm was successfully obtained. Residual stress was completelyrelieved by annealing treatment of 30min and permanent deformation was negligible.

A plate of 2mm was successfully obtained. Residual stress was completelyrelieved by annealing treatment of 30min and permanent deformation was negligible. Arelative density of 90%, which is standard level of commercially available part, wasobtained. Tensile and impact strength were limited to 1/2 and 2/3 of those obtained bynormal process, respectively. Energy consumption of laser module in preheating freeprocess is around 45MJ/kg, and complete robustness against power supply interruptionwas demonstrated.IntroductionTohoku Earthquake on March 11th 2011 and the following tsunami had entirelydamaged power generation infrastructure which was ranging more than 400km alongthe Japanese north east coast. As a result, The Tokyo Electric Power Company, whichnormally provides the greater Tokyo area with approximately 40,000MW at that time ofyear, suddenly lost 25% of its power generation.

As a result, The Tokyo Electric Power Company, whichnormally provides the greater Tokyo area with approximately 40,000MW at that time ofyear, suddenly lost 25% of its power generation. To prevent a sudden and followingchain reaction which might lead to entire blackout, the power company divided itsservice area into 5 sections and instituted daily rolling blackouts each of which lasts forthree hours. Implementation of the rolling blackouts had been expected to last for 1.5months, but, in reality, thanks to great effort for recovering power generation andsaving energy consumption, it had not been take place since one week after its initiationuntil today (July 15th 2011). Temperature rising in summer, requirement for electricityis increasing, and Japan is facing the fear of rolling blackout again. This issue is notlimited to the region which is directly struck by the quake or the tsunami, but involvingwhole Japan.

This issue is notlimited to the region which is directly struck by the quake or the tsunami, but involvingwhole Japan. After the quake, four nuclear reactors have been shutdown though theywere not damaged by the quake or the tsunami. Two shutdowns are due to the anxietythat another earthquake or tsunami destroys other nuclear plants, and the other twoare for routine inspection. Some reactors which were under inspection when the quakecaused were not rebooted even now though their inspection was finished already. As aresult, power companies request their customers to cut power consumption by 15% andannounce that they will take partial blackout if excessive demand is expected, to avoiddisastrous entire blackout.With respect to plastic laser sintering, power shortage matters for two reasons. First,large power consumption of the process itself is problematic.

First,large power consumption of the process itself is problematic. Figure 1 depicts powerconsumption of modules installed in a commercially available plastic laser sinteringmachine (Semplice, ASPECT Inc., Tokyo JAPAN) in idling state after warm-up isfinished. The machine has a relatively large build envelope of 550mm x 550mm. Asshown here, the machine consumes approximately 4.2kW during idling, and more than60% is spent on heaters which are controlling the temperature of powder.Cooling fans, 190W, 4%Controllers, 110W, 3%Scanners, 40W, 1%Other, 80W, 2%Servomotors, 30W, 0.7%Heaters, 2600W, 62%Water cooler, 1140W, 27%Laser, 20W, 0.5%Figure 1 Power consumption of a laser sintering machine in idling state after warmup is finished.For the second reason, interruption of electric service likely to happen in such tightenergy condition, limits the maximum size that can be obtained by the process.

In theprocess, the powder bed temperature is elevated between the crystallization andmelting points of the powder, to reduce the generation of thermal stress [1].Temperature drop during building process by some reason such as blackout results inirreparable part wapage, and the process resumption is impossible due to the collisionbetween the powder supplying device (roller or blade) and the warping parts even afterthe temperature is recovered. Thus maximum size that we can obtain from plastic lasersintering is as much as we can finish while electricity service continues. In reality, manyservice bureaus in the greater Tokyo area were forced to stop providing large sized partsthough their facilities had not been damaged by the quake or the tsunami.In plastic laser sintering, powder bed preheating is performed to reduce generation ofthermal stress by keeping the surface above crystallization temperature.

On the otherhand, in commercialized metal laser sintering or melting, preheating up to near themelting temperature is not implemented. Parts is fixed to a base plate not to warp [2]during sintering or melting process, and the residual stress is usually relieved by stressrelieve annealing after the process.The present research investigates feasibility of plastic laser sintering which does notuse preheating but fixes the parts to a base plate. Reducing the power consumption forthe preheating is expected to contribute to power saving dramatically. Additionally,absence of preheating might allow us interruption of laser sintering process. If we canrestart part building after shutting down of the machine, robustness of the fabricationprocess is drastically improved.Density of an obtained part is the primary index for its mechanical strength.

The set ofparameters that maximizes the density was searched for in experimental-based manner.In metal laser melting (sintering) processes, a supplemental and sacrificial structureknown as anchor or support is used to fix the parts when the obtained parts should beseparated from the base plate, or they have undercuts. These structures, if it is required,must be strong enough to resist the warping force, but, on the other hand, it should beas thin as possible for ease of following removal work [3]. Number of the supportsshould be as small as possible for the same reason, but large interval between supportsresults in local warpage. Design of support structure was optimized in terms ofthickness and interval. The effects of stress relieve annealing were evaluated.

Design of support structure was optimized in terms ofthickness and interval. The effects of stress relieve annealing were evaluated. Densityand mechanical strengths of the parts were measured and compared with thoseobtained by normal laser sintering.MATERIAL AND METHODSPOWDER, MACHINE, BASE PLATE AND PROCESS PARAMETERSA commercially available polyamide powder (DuraForm PA, produced by 3D Systems)was used. A machine developed by the authors [4] was used. The machine is equippedwith CO2 laser source (GEM-30, Coherent Inc.), which can emit a single mode laser witha wave length of 10.6µm at the maximum output power of 30W. The laser beam isfocused into a 130µm diameter (1/e2) spot on the powder bed with a 109mm focallengthed ZnSe lens. The beam can be scanned by a galvanometer mirror system in arange of 100mm×100mm. For powder supply, a roller system is used. The roller is20mm in diameter and rotates with a DC servo motor.

The beam can be scanned by a galvanometer mirror system in arange of 100mm×100mm. For powder supply, a roller system is used. The roller is20mm in diameter and rotates with a DC servo motor. As this roller system is placed ona linear actuator which is driven by a stepper motor, rotation speed and direction of theroller can be controlled independently to traversing speed. In this research, typicalcounter roller mode was used.The base plate and support structure should adhere tightly enough to resist strongwarping force, which is generated when melt plastic is crystalized and cooled to roomtemperature. For this purpose, the base plate was made from the same material, i.e.DuraForm PA, by laser sintering.

For this purpose, the base plate was made from the same material, i.e.DuraForm PA, by laser sintering. To prevent the plate from bending, the plate wasattached to a 6mm thick aluminum alloy backup plate and fixed with screws asdisplayed in figure 2.In the following experiments, laser power, scanning speed, scan spacing was varied tooptimize the parameter and to investigate the relationship between the parameters andthe obtained performance. Layering thickness and roller traversing speed and rotatingspeed were set at 100μm, 50mm/s and 150rpm, respectively.30DuraFormPA43 1075Aluminum alloy50[mm]Figure 2 Base plate structurePART DENSITY OPTIMIZATIONPart density optimization was performed by building coupons with various processparameters and comparing obtained densities. As the design of coupon, a thin plates of20mm×30mm×2mm was used. To process this coupon, 19 layers of laser scanningwere performed.

As the design of coupon, a thin plates of20mm×30mm×2mm was used. To process this coupon, 19 layers of laser scanningwere performed. Thus, it was always thicker than 18mm, and actual thicknesses werevarying between 19mm and 20mm. We use relative density (cid:2025)(cid:3045), which is calculated as: (cid:1865) (cid:1874)(cid:3028) (cid:3400) (cid:2025)(cid:3046)(cid:2025)(cid:3045) (cid:3404)(1)where (cid:1865), (cid:1874)(cid:3028) and (cid:2025)(cid:3046) are mass of a coupon, actual volume of the coupon and specific gravity of the powder material (0.97g/cm3 for DuraForm PA).OPTIMIZATION OF SUPPORT STRUCTUREA 1mm high mesh type support structure as illustrated by figure 3a was used. Tooptimize the width of the support, ws, overhang and support model as displayed infigure 3b was used and various support thickness were tested.

Tooptimize the width of the support, ws, overhang and support model as displayed infigure 3b was used and various support thickness were tested. Various gaps betweensupports, gs, were tested, and warpage between neighboring ribs was observed to findsupport density requirements.wsgssupport1mmws150μmabFigure 3 Mesh type support structure and its parameter optimizationRESIDUAL WARPAGE AND EFFECT OF ANNEALINGSupport can prevent undergoing parts from warping, but the parts might deform whenthey are removed from supports due to residual stress. Warpage by the residual stresswas measured by building test pieces of 50mm×10mm×2mm. Annealing at 80°C wascarried out for 30minutes and the effect was evaluated. For annealing, an industrialprogrammable oven, DO-450PA, was employed.

Annealing at 80°C wascarried out for 30minutes and the effect was evaluated. For annealing, an industrialprogrammable oven, DO-450PA, was employed. This oven is installed with 1.3kW ofelectric heater and possible to elevate the temperature ofits chamber of450mmx400mmx450mm up to 270°C.STRENGTH MEASUREMENTSTensile and impact tests (ISO 527 and 180, respectively) were carried out. Sincestandardized specimen for tensile test could not be obtained due to build envelopelimitation, specimen design as displayed in figure 4 was used. Build direction of eachspecimen was rectangular to the pulling direction in the tensile tests. The results werecompared with those data of normal preheating process. For fabrication of normalprocessed parts, a prototype machine (beta version of Semplice produced by ASPECTInc.) was employed. The process parameters were tuned so that the relative densitybecomes 90%, which is equivalent to a density value obtained without preheating.

The process parameters were tuned so that the relative densitybecomes 90%, which is equivalent to a density value obtained without preheating. Theparameters for the normal process with preheating are summarized in Table 1.9070542610Build direction48[mm]Figure 4 Tensile test pieceTable 1 Parameters for process with preheatingBeam power at the surface of powder bed Scan Speed Scan Spacing Powder Bed Temperature Density Tensile Strength Impact Strength Unnoched Izod13 W 3.78 m/s 150 µm 177 °C 89.70% 43.3 Mpa 24.2kJ/m2BLACKOUT SIMULATION TESTConfirm the advantage of laser sintering without preheating, a test as following wascarried out.1. On a 1mm high support mesh, a 20mm×30mm square shape coupon is built up tothe thickness of 2mm (19layers).2. Process is stopped and kept suspended for 30minutes.3. Restart the process to finish another 20layers.4.

Process is stopped and kept suspended for 30minutes.3. Restart the process to finish another 20layers.4. Defect on the interruption is observed visually.ENERGY CONSUMPTION MEASUREMENTAs a result of quitting preheating, laser must supply more energy than when preheatingwas performed. This increase should be smaller than energy consumption to let thepreheating free process be energy saving. Power consumption of the laser system ateach laser power was measured by clamping its power line with a current tester (HIOKI3293-50), and energy consumption was calculated by simply multiplying the read of thetester by 100V, outlet voltage in Japan.RESULTSPART DENSITY OPTIMIZATIONRelative density of typical laser sintered plastic parts is above 90%. In this research,parameters that can provide an equivalent density were searched for. Laser scanningspeeds, (cid:1874)(cid:3013), and scan pitches, (cid:1868)(cid:3046), as listed in table 2 were used.

Laser scanningspeeds, (cid:1874)(cid:3013), and scan pitches, (cid:1868)(cid:3046), as listed in table 2 were used. Laser power at the powder surface, (cid:1842)(cid:3013), was varied up to the maximum value of 17W. Figure 5 depicts relationship between relative density (cid:2025)(cid:3045) of energy irradiation per unit area (cid:1857)(cid:3028), which can be calculated as:(cid:1857)(cid:3028) (cid:3404)(cid:1842)(cid:3013) (cid:1874)(cid:3013) (cid:3400) (cid:1868)(cid:3046)As supplied energy density increases, density of part increases as general tendency. The maximum density of 93% was obtained when (cid:1857)(cid:3028) (cid:3404) 203kJ/m(cid:2870). Supplying large energy colored the parts and made them brittle.

The maximum density of 93% was obtained when (cid:1857)(cid:3028) (cid:3404) 203kJ/m(cid:2870). Supplying large energy colored the parts and made them brittle. A parameter set of (cid:1842)(cid:3013)=7.5W, (cid:1874)(cid:3013) (cid:3404) 1.5m/s and (cid:1868)(cid:3046) (cid:3404) 30μm, which brings a density of 90.9%, was selected as the default parameter set and used in the following discussion unless otherwise noted.Table 2 Parameters for part density optimizationScan speeds [m/s]0.92, 1.53, 3.81scan spacings [μm]20, 30, 40, 5010095]%[y t i s n e d9085e v i t a l e R80750.93 m/s 1.52 m/s 3.81 m/s70100150 Energy density [kJ/m2]200250Figure 5 Relationship between relative density and energy irradiation per unit areaOPTIMIZATION OF SUPPORT STRUCTURETable 3 summarizes characteristics of supports with various thicknesses. Nominalthickness is the distance between the center lines of the outermost scans.

Nominalthickness is the distance between the center lines of the outermost scans. Though thedistance between the center lines of scans is discretized by scan spacing (30μm as(2)default), we neglect the difference between nominal thickness and real distance ofoutermost scans to simplify our discussion. Additionally, actual thickness is thickerthan nominal thickness due to excessive sinter as listed in table 3.Support with nominal thickness of 300μm was strong enough to support fine structureat the middle of the parts but did not stand curling force at the edge as shown in figure6b. Secure support was obtained when its thickness was 0.4mm or thicker.Small warpage between supports is likely to occur along the edge of the coupon. Itstarted to occur around the test piece when the gap became larger than 2.0mm, andintruded into the middle section when it is larger than 3.0mm.

Itstarted to occur around the test piece when the gap became larger than 2.0mm, andintruded into the middle section when it is larger than 3.0mm. For ease of supportremoval, default gap between supports was set at 2.5mm and an extra support wasplaced along the outline of the part.Table 3 Strength of supportthickness [μm] actualstrengthusagenominal 100 200 300 400 500impossible to buildNA NA390 tear off easily 530 able to be tore off by pulling with tweezers fine structure at the middle 610 strong 700 very stronggeneral outlinea broken support at the edgeb warpage between supportFigure 6 Failure of supportFigure 7 Warpage after buildFigure 8 Warpages after removing from the base plateRESIDUAL WARPAGE AND EFFECT OF STRESS RELIEVE ANNEALINGFigure 7 shows typical appearance of warpage after a part is removed from its baseplate.

Figure 8 compares warpage when removed as soon as the sinter process isfinished, 30 minutes after the sinter process and after 30min annealing at 80°C. Asshown here, annealing treatment is a quite effective countermeasure against residualdeformation.STRENGTH MEASUREMENTTable 4 lists process parameters that were used in the strength measurements. Figure 9shows ultimate tensile strength for the process conditions. As shown here, tensilestrengths are around 21MPa, and difference between conditions is very rare. As anoverall tendency, the tensile strengths obtained by preheating free laser sintering arethe half of those by the process with preheating using parameters (43MPa as showntable 1).Figure 10 shows impact strength (izod unnotched) for the same process parameters astensile tests. In these results, we can find clear superiority with condition 3, which isslower in scanning speed, narrower in scan spacing and lower in laser power.

In these results, we can find clear superiority with condition 3, which isslower in scanning speed, narrower in scan spacing and lower in laser power. Theimpact strength obtained by this parameter set, 16.7kJ/m2 in average, reaches 2/3 ofthose obtained by normal process with preheating (24kJ/m2 as shown in table 1).Table 4 Parameter sets used in strength testsCondition1 2 3 4Scann speed [m/s] 1.53 1.53 0.92 0.92Scan spacing [μm] 30 50 30 50Laser power [W] 7.5 10.8 4.9 7.5Energy density [kJ/m2] 163 141 178 16325] a P M[20h t g n e r t s15e l i s n e t10e t a m5i t lU01234ConditionFigure 9 Ultimate tensile strengths in various conditions25]2m / J k [20h t g n e r t s1510t c a p mI501234ConditionFigure 10 Impact strengths in various conditions.

Izod unnotched (ISO 180)BLACKOUT SIMULATION TESTFigure 11 shows a closed-up side view of a part whose process had been interrupted for30min to similate a blackout. As far as observing visually, we do not find defect ordifference in suspended layer..Figure 11 Closed-up side view of a part whose process was interrupted for 30min.Suspended layer is indexed by the arrow.500]W400[n o i t p m u s n o c300200r e w o P100005 10 Laser power at powder bed [W]1520Figure 12 Relationship between energy consumption and laser irradiation at powder bedENERGY CONSUMPTION MEASUREMENTFigure 12 is relationship between power consumption, Pc[W]and laser intensity on thepowder bed, PL[W].

This relationship can be approximated by a parabolic function as:(cid:1842)(cid:3004) (cid:3404) 0.556(cid:1842)(cid:3013)(cid:2870) (cid:3397) 12.9(cid:1842)(cid:3013) (cid:3397) 73.2Though power consumption for (cid:1842)(cid:3013) (cid:3404) 0 is 73.2W according to this formula, actual power consumption which was measured while the laser was idling was 23.4W.DISCUSSIONSMECHANICAL PERFORMANCEAchievement of the high relative density more than 90% demonstrates that powder is(3)successfully melted and a high flowablity of the molten plastic was obtained as well.Large deterioration in tensile strength despite the high relative density is an issue to beconsidered. Since the temperature of substrate, the previously sintered layer, is colderthan the case of preheating process, delamination between layers seems to occur moreeasily.

Since the temperature of substrate, the previously sintered layer, is colderthan the case of preheating process, delamination between layers seems to occur moreeasily. However, since the pulling direction is rectangular to lamination, delaminationdoes not seem to play dominant role in the degradation of the tensile strength in thedirection. There might be two possible reasons. One is difference in microstructure [5],and the other is degradation of material performance. Latter suspicion is supported bydeterioration in impact strength as general tendency. In addition, smaller decrease ofimpact strength when laser power is relatively low indexes that irradiation of highlyintense laser might play important role in low mechanical performance of lasersintering without preheating.WARPAGE PREVENTIONAs far as building thin objects such as the test piece in the present research, preventionof warp distortion by using support structure is successful and satisfying.

However,effectiveness for such a thick and large object that it takes more than a day to finish itsbuilding is not clear. To reveal the feasibility of such applications, more analytical studyand quantitative measurements of warpage and warping force is required.We can expect that creeping ability of plastic facilitates the fabrication of large objectwithout preheating than metallic materials.POWER SAVING ASPECTTable 5 summarizes the laser power at powder bed, PL, power consumption, Pc,,irradiation time per unit volume, (cid:2028)(cid:3013)(cid:3036), and energy consumption per unit volume, eLc.

This is comparable to the consumption of the laser.(4)(5)Table 5 Energy consumption example of processes with/without preheating.Power consumption of laser unit is obtained from laser power at powder badusing formula 3.Parameters Laser power at powder bed [W] Power consumption of laser unit[W] Scan speed [m/s] Scan interval [μm] Layer thickness [μm] Irradiation time per voluum [s/cm3] Energy consumption by laser per voluum [kJ/cm3] Energy consumption by laser per mass [MJ/kg]preheat free 7.5 201 1.5 30 100 22245 43preheat 13 335 3.78 150 100 186 6CONCLUSIONPlastic laser sintering using support structures instead of powder bed preheating wasproposed. It was shown that a plate with thickness of 2mm can be obtained, and partdensity met standard level of normal preheating process, 90%. Investigation ofapplicability to larger or thicker object requires further research. Residual stress can becompletely relieved by annealing treatment for agreeable duration.

Investigation ofapplicability to larger or thicker object requires further research. Residual stress can becompletely relieved by annealing treatment for agreeable duration. Tensile and impactstrength were limited to 1/2 and 2/3 of those obtained by normal process, respectively.Clear reason for this degradation was not obtained. Energy consumption of lasermodule in preheating free process is around 45MJ/kg, and complete robustness againstpower supply interruption was demonstrated.REFERENCES[1] Beaman, et al, “Solid Freeform Fabrication: A New Direction in Manufacturing,” (1996)[2] Kruth, et al, “Feedback control of Selective Laser Melting.In” Virtual and rapid manufacturing(2008) pp 521–527[3] Jhabvada, et al, An innovative method to build support structures with a pulsed laser in theselective laser melting process,” Int J Adv Manuf Technol, published online 07 Jul.

ATURE EXTRACTION FROM TESSELLATED AND SLICED DATA IN LAYERED MANUFACTURINGKamesh Tata Prototype ExpressGeorges Fadel Clemson UniversityAbstractWhen parts are built in layers, the cross sectional area of each layer has to be de fined and filled with a pattern of vectors. This filling process is called hatching and the vectors defme the hatch pattern. To accurately reproduce a three dimensional object, In particular, top and bottom surfaces, edges. holes key features need to be identified. and protrusions must be recognized to ensure the slice plane does carry-the critical in formation required for the build. This paper describes a technique to extract relevant features from a tessellated model to generate a correct sliced representation.IntroductionLayered based manufactUring is based on a mathematical slicing operation that generates slices from an original boundary representation model (tessellated or triangu lated model).

Different processes use different devices to draw the hatch vectors \vhich represent the ro.;eas to be filled or solidified in a slice.

For instance, Stereolithographyand Selective Laser Sintering use a laser to draw the vectors, while other processes may use a pattern mask or a binding agent.Many slicing engines exist, in particular, the slicing program of 3D systems has a proven algorithm, Grogan [1990] described an algorithm based on sorting, Chalasani [1992] proposed two algorithms for slicing 3D objects that also rely on a sorting process to identify areas, Vouzelaud and Bagchi [1992} proposed an adaptive technique, Kirsch man and Jara-Almonte [1992] investigated parallel slicing, and Dolenc and lY!akela (1993) suggested slicing procedures which considered the identification of some features, notably peaks and flat areas.This paper considers the issue of features and their extraction from the tessellated It highlights the capabilities of such an algorithm to identifY manufactUring feadata.

tures and lists the pitfalls and their possible remedys.The STL FileThe de facto industry STL model representation defines CAD solid models as a set of triangular facets (3D Systems, 1989). Facet models represent solid objects by spatial boundaries which are defined by a set of planar faces. This is a special case of the more general boundary representation which does not require object boundaries to be planar (Mortenson, 1985). In general, the term facet is used to denote any constrained polygonal planar region used to define a model boundary; however, in the SFF community, the term facet is typically understood to mean triangular facet (Rock. 1991). Unfortunately.

1991). Unfortunately. these facets are stored independently, as if each facet was created and tossed into a bucket with no particular ordering and without information relating a given facet to any other facet in the bucket (Rock, 1991).Hole or ProjectionOne of the time consuming tasks of any hatching routine is identifying holes and projections. Holes should not be hatched. while projections should. At the time oil1atch-587ing, the information about the solid is no more available in a single piece. Most of it is divided and stored in a number of 2-D slices and some of it is lost. This implies that iden tifying a single hole or a projection might require studying several slices together. This is computationally expensive.

This implies that iden tifying a single hole or a projection might require studying several slices together. This is computationally expensive. Thus there is a need for an efficient hatching algorithm which identifies, in the qUickest possible way, areas that need to be filled and areas that need not be filled..Strategy for HatchingIn order to identify the areas that need to be hatched, we examined various contours and derived the following rule: If a contour is surrounded by n contours and n is an even number, the contour represents a protrusion and should be filled. On the other hand, if n is odd, it the contour represents a hole and should not be filled. This rule, \vhich is central to the hatching algorithm developed as part of this research, can also be explained in another way.

This rule, \vhich is central to the hatching algorithm developed as part of this research, can also be explained in another way. Considering 3-D solid objects,1.a solid is always bounded,2.there cannot be a solid without thickness,3.and there cannot be a hole without a solid.These three simple and definitive statements ensure that when a solid is sliced the resulting contours are inVariably closed and non-intersecting. Also, there will never be two holes adjacent to each other or one contained in the other without being separated by a definite distance. Furthermore, any unbounded contour (contours not surrounded by any other contour) can never be a hole.Key Characteristics IdentifierA simple knowledge base, supported by a few rules, is created with an aim to recognize the following from a tesselated model:1. base faces of features,2.t:Y--pe of features (protrusion or depression), and3.

base faces of features,2.t:Y--pe of features (protrusion or depression), and3. geometric shape and orientation of features.The rules are mostly based on simple geometric and engineering principles, and derived by observation and logical reasoning. Some of the ideas discussed can be utilized for any solid model regardless of the representation scheme.The Three Types of Base FacesThe top and bottom faces of any feature. including the object. may be termed as base faces. Recognition of base faces is essential to improve the tolerances of layered models. Since a polygon is comprised of only two basic entities, point and line. only their combina tions will result in any topologically significant characteristic. Consequently, any solid represented by planar polygons can have only three types of base faces:1. Horizontal surface: A horizontal surface is represented by a single or multiple polygons parallel to the slice plane.

Horizontal surface: A horizontal surface is represented by a single or multiple polygons parallel to the slice plane. Figure 1 shows horizontal faces composed of triangular facets.2. Pointed edge: When at least two non-co planar polygons share an edge, the edge may be termed as a pointed edge. This is illustrated in Figure 2, where a block with v-grooves is shown.588Figure 1. Multiple Facets Forming Horizontal and Vertical SurfacesPointed EdgesFigure 2. Block with Pointed Edges3. Pointed End: When at least two non-eo planar polygons share a venex, the vertex may be termed as a pointed end. Tip of a cone and corner of a block, shown in Figure 3, are examples of pointed ends.Figure 3. Examples of Pointed EndsFor the purpose of layered manufacturtng, the effort must be to identify the exis tence and then to determine the height (along the slice axis and from a reference point) of each one of the above so that the slice plane can be forced to assume that height.

Identi fication of these three types from an STL model format is basically a searching operation.589Identifying Horizontal SurfacesHorizontal surfaces can be identified by checking unit normals or ordinate values of facets. If the unit normal of a facet describes 90° (or 270°) with the slice plane, there lies a horizontal surface with that facet contained in that surface. Alt-ernatively, if the z coordinate values of all the three vertices of a facet are equal, then that facet is part of a horizontal surface.Vertical walls can be identified essentially using the same apRroach used to identify horizontal surfaces. lfthe unit normal of a facet makes 00 (or 1800) with the slice plane, there exists a vertical wall with that facet contained in that wall.Identifying Pointed EdgesAccording to the definition of a pointed edge, any edge shared by two non-co planar facets is a pointed edge.

With this definition, there are twelve pointed edges, four vertical and eight horizontal, in a cube shown in Figure 4. It does not serve any purpose in the context of layered manufacturing to identify these edges separately, since me vertical and horizontal surfaces of which these edges are part of, can be easily identified using meth ods suggested earlier. Likewise, it is worthless to identifY the edges of a tetrahedron shown in Figure 4 because none of the edges of the inclined surfaces are parallel to the slice plane. As we discussed previously, our primary goal is to identifY key characteristics which are parallel to the slice plane so that we can retain them in the sliced model .Figure 4. Unwanted Pointed EdgesIt is, therefore, necessary to narrow the definition of a pointed edge, in the context of layered manufactUring. First, none of the polygons sharing an edge should be horizon tal or vertical.

First, none of the polygons sharing an edge should be horizon tal or vertical. This is due to that horizontal and vertical faces are already identified and with that all their edges. Further, the common edge must be parallel to "the slice plane. (Only then a separate effort is required for its retention in the sliced model.)A pointed edge can now be defined as: Any non-horizontal and non-vertical facet with two vertices having the same z coordinate value will form a pointed edge provided the second facet that shares these vertices (there will be only one such facet) is not in the same plane as the fIrst facet and is non-horizontal and non-vertical. Also, the angle be tween those two facets need to be defmed by the user. This is to avoid mistakenly identify ing the edges of facets approximating curved surfaces as pointed edges. If we consider the faceted model of a sphere, according to the above definition, potentially every edge becomes a pointed edge.

If we consider the faceted model of a sphere, according to the above definition, potentially every edge becomes a pointed edge. This can be avoided by defining the angle between the two facets sharing an edge to be less than a certain value, say, 120°, before the edge can be consid ered as a pointed edge.This new and narrow definition of a pointed edge helps reduce the time required to search for such edges in a faceted model. However, identifying two vertices of a facet with the same z coordinate value, the very first step in identifYing a pointed edge, demands virtually every vertex of all non-horizontal and non-vertical facets in the model to be searched. This again leads to high execution times.590Identifying Pointed EndsFinally, pointed ends can beJdentified from an STL flie by looking for a single vertex shared bv several non-co planar facets. For example, the apex of a cone is shared by a number of non-co planar facets which form the conical surface.

For example, the apex of a cone is shared by a number of non-co planar facets which form the conical surface. This calls for checking everY vertex in the model for the number of facets that share the vertex and their unit normals.Base Faces - Change in the Number of ContoursClearly, identifying pointed ends and edges is a time intensive search operation. Isolation of a vertex that forms a pointed end or a facet side that forms a pointed edge demands a large scale search often covering every facet of the model. This problem can be largely overcome by adopting a new strategy. This strategy is based on the fact that whenever there is a change in the number of contours in successive slices, there lies a base face of a feature. The exact location of the base face can be anywhere between the two slice planes or contained in the slice plane which has the greater number of contours.

The exact location of the base face can be anywhere between the two slice planes or contained in the slice plane which has the greater number of contours. Once a change in the number of contours is observed, the exact location of the base face can be found by considering facet groups in that region and applying the methods sug gested above. However, this strategy requires slicing the model and counting the number of closed contours in each slice.Slice Based Feature Recognition - Protrusions or DepressionsProtrusions are any projections on the object, such as pads. bosses, and mounds. while depressions are holes of any shape. The easiest 'way of finding the type of a feature is by studying the slices and the contours within each slice. The following rules will pro vide methods to distinguish between protrusions and depressions.

The following rules will pro vide methods to distinguish between protrusions and depressions. These rules are sim ple, definite, unambiguous, and true for any solid.Rule 1: When a solid is sliced by an imaginary plane, the resulting slice may contain a single or multiple, non-intersecting, closed contours. This is true regardless of the angle and the position of the slice.Before going further, it is necessary to define certain key words used in the preceding part of this section.1. Empty contour: An empty contour is one which does not contain any other con tour within its bounds. In other words, an empty contour does not surround any other contour. See Figure 5.2. Unbounded contour: An unbounded contour is one which is not surrounded or bounded by any other contour.Rule 2: If a single or multiple empty contours are surrounded by a single larger contour, all inner contours represent depression features.

This can be seen in Figure 6, where a rectangular block \vith several holes is shown. The slice at AA has several empty contours surrounded by a single large contour. Obviously, all inner contours represent holes and hence the rule 2 is satisfied.Rule 3: If two or more contours are nested within each other, all even-numbered contours represent depressions while all odd-numbered contours represent protrusions, the outermost contour being number 1. Figure 7, where a cylindrical block with multiple holes one inside the other is shown, illustrates rule 3.Rule 4: Any unbounded contour represents a protrusion feature.Figure 8, where a number of protrusions are shown on the top of a rectangular block,591Unbounded contourEmpty contourFigure 5. Empty and Unbounded Contours,A~ __ II - -,-, , I,-I " i - - . - - .,.--- - , - - , - II1,l._ _ ,"I ,\",.JA°Dt60Front view of a block with four holesSlice at AA (view along z)Figure 6.

- - .,.--- - , - - , - II1,l._ _ ,"I ,\",.JA°Dt60Front view of a block with four holesSlice at AA (view along z)Figure 6. Multiple Empty Contours within a Larger Contourclearly illustrates this point. The slice at AA contains several unbounded contours and all are, obviously, protrusions. This satisfies rule 4.Shape and Orientation of FeaturesOnce, location and type of a feature are determined, further details can be obtained by studying the unit normals of the facets the feature is comprised of. The following as sumes that holes and projections. and base faces are already identified using the meth ods suggested above.A feature is rectangularly shaped, if it is bounded by four faces, and each face is perpendicular to its adjacent face. Again, all four must be perpendicular to a con1mon plane. If a fifth face that connects all four faces exists, the feature is closed at one end.

Again, all four must be perpendicular to a con1mon plane. If a fifth face that connects all four faces exists, the feature is closed at one end. It is closed at both ends if a sixth face also exists connecting the first four faces. Further more, the feature can be a rectangular pad or a rectangular hole depending upon its na ture (protrusion or depression). Triangular pads or holes are bounded by three faces, where the sum of the included angles between the faces is exactly 1800 and all three are perpendicular to a common plane. The type of triangle can also be easily established by studying the included angle between faces. This logic can be extended to identify penrag-592I II IA I ___ - - ' - - _ -.L..-J-III ...L....J-... _ ....J..-.- _IIAI I ...I I - _1- ...II. _I I I I .. _.I I "Front view of a cylinder with two holesSlice at AA (view along z) Grey indicates solid White indicates holeFigure 7.

_I I I I .. _.I I "Front view of a cylinder with two holesSlice at AA (view along z) Grey indicates solid White indicates holeFigure 7. Multiple Contours Surrounding Each OtherAA o DOSlice at AA (view along z)Front view of a block with three protrusionsFigure 8. Unbounded Contours593onal, hexagonal or any polygonal feature. Orientation of the feature can be established by determining the angle between the slice plane and one ofthe bounding faces. Different types of features that can be easily identified are shown in Figure 9.Rectangular PocketTriangular PocketOpen SlotFigure 9. Typical Features Which Can Be Identified from Tesselated ModelsOne limitation with slice based feature recognition is that features parallel or at cer In order for a comprehensive tain angles to the slice plane will be difficult to identifY. feature recognition, slicing may have to be done at different angles and the resulting slices studied independently and together.

feature recognition, slicing may have to be done at different angles and the resulting slices studied independently and together. For example, in Figure 10, holes A a.T1d B, which are parallel to the slice plane, cannot be identified. Hole C, which is perpendicular to the slice plane, can be easily identified as a depression feature. In order to identifY A and B, slicing has to be done along x ory.Feature recognition is a vast area and the 'work done here is only a small effort toward that goal.--E$c- -9---;B xVFigure 10. Block with Three Holes Along the Three AxesConclusionsThis work describes a possible methodology for the identification of features from layered or tessellated data. A simple rule base is used to identifY such features. The method does have significant limitations since features are only recognized when al0n.g the build axis.

A simple rule base is used to identifY such features. The method does have significant limitations since features are only recognized when al0n.g the build axis. However, if one considers what features are used for, especially to identitY hatching areas, then this feature recognition is sufficient. If the features are needed for594design purposes, then the approach has to be performed three times, one for each coordi nate axis as the build direction.References1.3D Systems Inc., (Valencia, CAl, 1989, "Stereolithography Interface Specification."2. Chalasani, K. L., Grogan, B., Bagchi. A., Jara-...~monte, C. C., Ogale, A. A., and Dooley, R. L., "An Algorithm to Slice 3-D Shapes for Reconstruction in Proto typing Systems", Computers in Engineering - Volume One, ASME 1991.3. Chalasani, K. L., "Design and Implementation of Slicing and Hatching Algorithms for Freeform Fabrication", M. S. Thesis, Clemson University, Clemson, S. C., 1992.4.

atures in Layered Manufacturing of Heterogeneous ObjectsXiaoping Qian xpqian@engin.umich.eduDebasish Dutta dutta@engin.umich.eduDept. of Mechanical Engineering and Applied Mechanics The University of Michigan Ann Arbor, MI 48109-2125ABSTRACTThe usage of features in computer aided design and manufacturing has increased significantlyover the last decade. By and large, all such features are geometric (form features).In this paper, we shall discuss the need to go beyond current geometric features and consider material -- composition, and gradation -- within the object. This need has been brought about by layered manufacturing technologies which build up parts, layer by layer, under computer control. While industrial use of this new technology has been for making prototypes, functional metallic parts can/are being made by layered manufacturing.

While industrial use of this new technology has been for making prototypes, functional metallic parts can/are being made by layered manufacturing. Furthermore, a variety of materials can be deposited to create multi-material and functionally graded components.We consider features in this new domain (of layered manufacturing) and identify research top ics and present an overview of our current focus on "material features" in the context of heteroge neous solid models.1.0 IntroductionLayered manufacturing(LM) refers to the host of manufacturing technologies that deposit material, layer by layer, under computer control. It is also referred to as solid freeform fabrication (SFF) and in industry as rapid prototyping. Layered manufacturing methods can be used to make functional parts which have good surface quality and good material property.

Layered manufacturing methods can be used to make functional parts which have good surface quality and good material property. One distinguishing characteristic of LM is that it is capable of depositing multiple materials in a layer as welLas vary ing the material composition. Harnessing this characteristic is the focus of our research.In this paper, we consider heterogeneous objects i.e., objects made of different constituent materials and having continuously varying material composition and/or microstructure (See Fig. 1). The state-of-the-art design and analysis research indicatesthatheterogeneous objects nat urally arise in the context of thermal expansion coefficients, mechanical strength, creep,. thermal and electrical conductivity, etc. Applications ofheterogeneous objects include automotive engi neering, aerospace, and biomaterials, etc.689(a) Cutting tool (diamond/SiclPolymer) (Y. Miyamoto, Ceramic Trans Vo1.76)(b) Cross-section from ALCOA injection molding tool( 1.

Miyamoto, Ceramic Trans Vo1.76)(b) Cross-section from ALCOA injection molding tool( 1. Fessler, et aI. SFF 1997)FIGURE 1. Heterogeneous objectsA key requirement for the fabrication of heterogeneous objects is their accurate representation in CAD systems. Solid modeling techniques for heterogeneous objects are just beginning to emerge [1][2]. However, there is a need for tools that can facilitate the fabrication process plan ning of such objects. Process planning for layered manufacturing can be a complicated task, involving the determination of orientation, slicing, deposition path planning, etc. In addition, fab rication of heterogeneous objects demands very delicate control of deposition to achieve desired material composition and gradient.In our research, we explore the use of features to facilitate the high level conceptualizing of material composition and gradation and its downstream transforming to layered manufacturing instructions.

Features were initially proposed to automate the link between design and NC path generation [3]. Since then, features have been widely and successfully used in CAD/CAM. Fea ture based design expedites the design process and feature recognition facilitates the fabrication process planning. A feature based product model also simplifies the assembly and inspection planning, etc. [4].In this paper, we report on our ongoing work. We propose a feature based architecture for design and process planning for layered manufacturing. In this context, we present several feature related issues. We then describe how we propose to address these issues and conclude with a sum mary.2.0 Feature-based design and process planning for LMOur proposed feature based design and fabrication integration approach is illustrated in Fig. 2.

2. In this scheme, design can be carried out either by a feature based design method (explicit design whereby material composition is explicitly imposed over the geometry) or by an implicit design method (whereby material information is derived such as in homogenization design method). In feature based design method, the design features are mapped to fabrication feature. In implicit design. method, a heterogeneous CAD model is employed to represent the design and a feature recognition module is adopted to extract the fabrication feature from this heterogeneous CAD model.690Feature MappingFeat~re based Designl-olll---1Orientation CalculationSlicingPath PlanningFIGURE 2. Feature based design and process planning for layered manufacturingUnderstanding and characterizing the layered manufacturing processes is essential to generate the correct fabrication strategy.

Feature based design and process planning for layered manufacturingUnderstanding and characterizing the layered manufacturing processes is essential to generate the correct fabrication strategy. Feature based process planning is carried .out on the basis of these characterization as well as the identified fabrication features. A set of mapping rules enable the identification of fabrication strategies, including orientation,slicing, road width, deposition path, etc., from the fabrication feature model. A final step might be post processing to generate the machine-specific data.3.0 Features related issues in the LM domain3.1 Design feature and fabrication featureDeveloping the concept of design feature in the context of heterogeneous object is of impor tancein the feature based design approach. In this context, a closer look at how heterogeneous object design is done will be helpful. Furthermore, methods representing the design feature is also an important issue.

In this context, a closer look at how heterogeneous object design is done will be helpful. Furthermore, methods representing the design feature is also an important issue. In addition, the interactions among features 9fdifferent ll1a.terial cOll1Positions demand specific feature operators beyond the conventional boolean operators. Finally, a library of design features can be developed to supportthe feature based designapptoach.Similarly, understanding and identifying fabrication features in the context of layered manU facturing processes is necessary for the automation of process planning. This fabrication feature identification will be at the core of the fabrication strategy generation. Also an efficient represen... tation of the fabrication feature is important.691Implicit design schemes such as the homogenization design method yield heterogeneous objects[5]. Creating heterogeneous solid models from such density data is a topic of ongoing research.

Creating heterogeneous solid models from such density data is a topic of ongoing research. For processing such designs, feature recognition algorithms will be necessary to extract the fabrication features from the heterogeneous solid model.3.2 Mapping from design feature to fabrication featureConsider the heterogeneous object shown in Fig. 3. Ouring its design (Fig. 3a), the groove would be··a design feature. However, for layered manufacturing, the same groove, becomes a fab rication feature made of (sacrificial) support material (Fig. 3b). As illustrated by this simple example, there is.a need to identify situations requiring design-to-fabrication feature mapping and develop efficient mapping methods.(a) Design(b) FabricationFIGURE 3. Mapping from design feature to fabrication feature3.3 Co-existence of NC feature and LM featureIt is increasingly becoming evident that in the future, manufacturing will include both NC machining and layered manufacturing.

Hence, it is important to develop methods that exploit the symbiotic relationship between NC and LM -- material removal and material deposition. Conse quently, feature extraction should be looked at within an enlarged context -- the synthesis of NC machining and layered manufacturing. Based on the extracted NC features and LM features, the optimal balance between NC machining and layered manufacturing has to be determined.4.0 Our current research4.1 Characterization of material variationHeterogeneous materials arises in shape and materially optimized structures. They can pro vide a smooth transition between materials which are otherwise incompatible because of their mechanical or chemical properties. They can also be used as a coating to modify the electrical, thermal, chemical or optical properties of a substrate [6].

They can also be used as a coating to modify the electrical, thermal, chemical or optical properties of a substrate [6]. The material variation is usually along some specific direction, based on the functionality and design intent and can be explicitly cap tured by a swept volume representation (See Fig. 4).692FIGURE 4. Swept Material VolumeSwept Material Volume [Working Definition]: Swept Material Volume is composed of a cross-section and a path. The material could vary along the cross-section and/or along the path.Conventionally, a swept volume S (Fig. 4) is defined by sweeping a surface r(u,v) along pathpew). If '*' denotes sweep, the swept material primitive (Fig. 4) could be written asSM(S, M) = ( r(u,v) * p(w), m(u,v,w) )where S represents the geometric volume, M represents the material composition, and mate rial variation function m(u,v,w) represents the material changes along the cross section r(u,v) and path pew).

The material variation function could be any user defined function (e.g., constant, lin ear, step function, parabolic, exponential, etc.).In practice, material variation typically would happen only along one of the u, v, w directions, denoted as m(u), m(v) and mew). The function m(ll,v,w) would enable full three dimensional material variation.This characterization of material variation provides a simple way to impose and manipulate material composition over geometry. The actual part could be composed of a combination of these swept material volumes. Similar to a CSG object in current solid modeling, material surVival rules in the synthesis of such volumes is an issue of ongoing research.4.2 Design featureThe design feature volume is composed of a swept material volume.

The following (working)definition extends the conventional feature by adding the material dimension:Heterogeneous Design Feature= (Conventional) Form Feature + Material CompositionNariation6932FIGURE 5. Design Featnres for Heterogeneons ObjectIII the sample part shown in Fig. 5, there are four design features (numbered 1 throl.lgh 4). Each feature has a .geometry and its own material composition (and fulfills a desi~n function). Feature 1 is the base of the part and made of homogeneous material. Depression feature 2 is a groove. Feature 3 is made of allother homogeneous material that is different from the base part material. Feature 4 provides the transition between 1 and 3 and makes the material property (e.g. coefficient of thermal expansion) transit smoothly from feature 3 to feature 1.

Feature 4 provides the transition between 1 and 3 and makes the material property (e.g. coefficient of thermal expansion) transit smoothly from feature 3 to feature 1. In this way, feature based design can provide the designer an efficient way to synthesize heterogeneous objects.4.3 Fabrication featureLayered manufacturing parameters for heterogeneous objects would consist of layer thick ness,foad width, material composition and its gradation. Based on this characterization,fabrica tion feature can be thought as a subvolume that can be produced by LM in the same deposition configuration (Le., material composition and gradation, road width, and layer thickness, are kept "constant"). We have the following (working) definition:LM Fabrication Feature= Subvolume + Deposition ConfigurationThe fabrication feature is a subvolume of a part that is distinguished either by the material type, material gradation, or the fabrication method.

The vertical.colllpoueut oflllaterial gradieut in ouesliceFig. 7 illustrates how material composition affects the slice thickness and road width.(a) G=O(b)(c)G...G, (d)(e)S: slice direction G: gradientFIGURE 7. Material Gradation Direction versus Slice DirectionDenote the slice direction as S, the gradient as G, e is the angle between S and G,o is theallowed material compositi?n tolerance. Then, the follo\\ling rules can be used... .. ..If G=O (Fig. 7a), it is a homogeneous object and the material composition can be ignored. If GII S (Fig. 7b), the slice thickness can be 0/0, and the road width can be maximum. If G.lS (Fig. 7c), the slice thickness can be maximum, the road width can be o/G.695If e is the angle between Sand G (Fig. 7d, e), the slice thickness can be o/Gcose, and the road width can be 0/G sin e .5.0 SummaryIn this paper, we described feature related issues in the layered manufacturing of heteroge neous objects.

An architecture for a feature based CAD/CAM system for layered manufacturing was presented. We proposed working definitions of swept material volume, design feature and fabrication feature. We also provided an overview of our ongoing work.6.0 AcknowledgmentThe authors gratefully acknowledge the support provided by ONR grant N00014-97-1-0245and NSF grant MIP 9714951.7.0 References1. Kumar, V. and Dutta, D., "An Approach to Modeling and Representation of Heterogeneous Objects", accepted for publication in ASME Journal ofMechanical Design, available as techni cal report, 1998.2. Marsan A. and Dutta D., "On the Application of Tensor Product Solids in Heterogeneous Solid Modeling", 1998 ASMEDesign Automation Conference, Atlanta, GA, Sep. 1998.3. Grayer, A., "A Computer Link between Design and Manufacture", Ph.D. Dissertation, Univer sity of Cambridge, UK, Sep.•·1976.4. Shah, J. 1., Mantyla, M., and Nau, D.

edforward Control for Polymer Laser Sintering Process Using Part GeometryMostafa Abdelrahman* and Thomas L. Starr**Rapid Prototyping Center, **Chemical Engineering Department, J.B. Speed School of Engineering, University of Louisville, Louisville, KY 40292AbstractFor the polymer laser sintering process, achieving optimum mechanical properties requires that every volume element of a part experience a temperature history sufficient to reach full density. This history must include a peak temperature high enough to fully melt, but not degrade, the polymer and a cool-down period that ensures elimination of porosity, interlayer bonding and relaxation of stress. Real-time thermal monitoring of the laser sintering process has shown that this temperature history depends on the geometries of both the current and prior layers. In this paper we demonstrate a feed-forward control system that improves uniformity of the temperature history for parts with variable cross-sections.

In this paper we demonstrate a feed-forward control system that improves uniformity of the temperature history for parts with variable cross-sections. The control algorithm for this system will utilize information from layerwise geometry models for parts in a multi-part build. The cross-sectional area for every layer will be used at run-time for feed forward control the laser scan parameters. The results confirmed maintaining constant peak temperature throughout the part. This control system ensures optimized sintering for parts with complex geometries.IntroductionPolymer laser sintering is an additive manufacturing method that converts layers of polymer powder into a dense 3D structure by selectively heating areas with a scanning laser. While this technology was developed more than 30 years ago all commercial machines operate with open loop control and limited monitoring and control of the process. Temperature control is a critical problem for the sintering of polymer powders.

Temperature control is a critical problem for the sintering of polymer powders. Achieving optimum mechanical properties requires that every volume element of a part experience a temperature history sufficient to reach full density. Process monitoring and control can be used to certify that a process is running consistently and within specifications for a particular material and process. A set of parameters can be determined to be controlled during the fabrication process. Examples of these parameters include the powder bed temperature and the peak temperature during the scan. Controlling this parameter set can establish the range of parameter values that produce "good" parts. Information from layerwise geometry models for parts can be used to control the process parameters. These geometric information can be used to adjust the laser beam power and other scan parameters.Previous investigators used optical sensors for laser sintering process monitoring, [1, 2, 3, 4,5,6].

All of this work analyses the optical emission results from the melt process in one or more spectral ranges. Diller et. al. [7] proposed to use thermal imaging for monitoring the laser sintering process. They proposed to analyse the temperature distribution on the part bed surface. Thermal imaging systems have been integrated into the Laser Sintering systems to monitor the thermal history, correlate the outcomes, and plan an optimized approach for part manufacture in [8, 9, 10,11]. However, these methods do not provide a quality certification of the process, and do not suggest how to use the monitor data to provide control signals for the sintering process.A new control method that takes advantage of the unique characteristics of the layerwise AM processes is needed. This method will use process control techniques to maintain the optimum process parameters.

This method will use process control techniques to maintain the optimum process parameters. These control signals will maintain a uniform temperature distribution of the powder bed surface in order to achieve equal part quality. The control signals may depend on the geometry of the layers and the process history of previous layers. One of the most important parameters that can be used to control the process is the cross section area of the layers.

One of the most important parameters that can be used to control the process is the cross section area of the layers. Our previuos work about the 3D indexing [12], with the proposed control signals will enable: setting up optimal laser beam parameters prior to the manufacturing process, controlling laser beam parameters during the manufacturing process, controlling the wait time before adding powder for the next layer, inspecting the quality of the current layer online, and creating a 3D Quality Certificate for the part being built.Process Control SystemsProcess control systems can be classified generally as open-loop, feed-forward or feed- back (closed-loop). Current commercial polymer LS equipment employs some elements of each control method. Normally a “set point” for laser power is determined prior to the start of the build and this output is maintained throughout the build period. This open-loop control corresponds to no “sensor” or “disturbance” inputs in Figure 1.

This open-loop control corresponds to no “sensor” or “disturbance” inputs in Figure 1. Although rarely used, current LS software from 3D Systems does have a provision for changing laser power as a function of build height. Build height would be a “disturbance” in Figure 1 and it would feed forward to the change the output laser power. The powder bed heater subsystem in current LS machines employs a simple closed- loop control method. A pyrometer measures the temperature at a point on the powder bed surface. This “sensor” signal is fed back and compared to the “setpoint” value and the difference is use to control the heater power. By design this system can only control temperature at one point on powder bed surface and has no provision for incorporating feed-forward information about layer geometry.The shortcomings of current commercial LS control systems lead to unacceptable variability in material properties, as discussed above.

The improved polymer LS process control system proposed here will include elements both of open and closed loop process control. Scan and temperature parameters will be set prior to the build but can be adjusted layer-by-layer based on the part and packing geometry of the layer (disturbance) and on real-time temperature measurements (sensor). Layerwise thermal imaging data will provide a richer set of monitor information that can be fed back to the control system. Control models (G’s) will be designed in future work to convert these inputs into control signals.Figure 1. Process control can be achieved using a combination of feedforward and feedback methods. Models (G’s) are used to convert inputs into control signals.In our previous work in [14], we show that there is a relationship between the peak temperature value and the geometry of the part, the layer cross-sectional area, and the presence or absence of an underlying solid layer.

We aproved that control signals can be generated to control the fabrication of the next layer and the wait time before adding the powder for the next layer. And we show that the cross section area can be used as a control signal to control the laser scan parameters. In this paper we propose to use the layer cross-sectional area to control the laser beam power in order to maintain a uniform peak temperature inside the build volume.Data CollectionFig. 3 shows the experiment setup. The system consists of a new open architecture polymer LS machine, thermal camera, 2 in flat gold mirror, and a computer. The process is monitored by a FLIR SC7000 MWIR thermal camera. The thermal camera is mounted as shown in Fig. 3 to view the building bed through the top laser window. The laser window and the mirror have a transmittance equal to 53% in the camera’s wavelenth range. The camera is focused manually on the region where the part will be built.

The laser window and the mirror have a transmittance equal to 53% in the camera’s wavelenth range. The camera is focused manually on the region where the part will be built. For this experiment, a 50 mm focal length lens is used and the frame rate set to 100 fps with window size of 255x175. The SLS machine takes about 230 minutes to build the one of the test parts. One SAF video is recorded for every part build. Every recorded video size is about 100 GB.Data AnalysisThe raw intensity data is mapped to the real temperature values using a 1st-order mapping function. The transmittance is set to 53% to compensate for the laser window and the mirror attenuation. The frames are parsed into individual layers using the using the maximum temperature in each frame as a marker as described in [12]. We selected 100 frame during the scan time for every layer. For each set of 100 frames the peak temperature value is detected for each point in the window area.

We selected 100 frame during the scan time for every layer. For each set of 100 frames the peak temperature value is detected for each point in the window area. These peak temperature values are assembled to form a single image for each layer. These images will be used to studey the effect of changing the laser beam power as a function in the cross-sectional area.Figure 2. System setup consists of the new “open architecture” LS machine, thermalcamera and a 2 in flat gold mirror.Template to Image RegistrationThe template layer (source) are registered to the corresponding peak temperature image (target) by estimating a projective transformation between the source template and the target image. The transformation is estimated between the edges of the template and the edges of the initially segmented image using fixed threshold.

The transformation is estimated between the edges of the template and the edges of the initially segmented image using fixed threshold. Given a source shape model 𝐌𝐬 that represents the template layer edges, and a target shape 𝐌𝐭 that represents the initially segmented image edges, a transformation 𝐀 that moves points from 𝐌𝐬 to 𝐌𝐭 is needed. The selected transformation has a set of parameters that will be estimated to minimize a certain energy function, similar techniques used before to estimate an affine transformation. Similar to [12], assume that the transformation is a 2D projective and hence it will have the following homogeneous format: 𝑎1 𝑎4 𝑎7𝑎2 𝑎5 𝑎8𝑎3 𝑎6 𝑎9𝐀 = ()(1)Assume that 𝑥𝑖 ∈ 𝐌𝐬, 𝑖 = 1,2, . . . , 𝑁𝑠 and 𝑦𝑖 ∈ 𝐌𝐭 is the closest point to 𝐀𝑥𝑖 on the target shape𝑥𝑖1 𝑥𝑖2 1where 𝑁𝑠 is the number of points on the source surface. Note that both 𝑥𝑖 = () and 𝑦𝑖 =(𝑦𝑖1 𝑦𝑖2 1) points are put in the homogeneous vector notation of size 3 × 1.

Note that both 𝑥𝑖 = () and 𝑦𝑖 =(𝑦𝑖1 𝑦𝑖2 1) points are put in the homogeneous vector notation of size 3 × 1. Consider the Euclideandistance between the moved point and its closes position to be the dissimilarity measure as follows: (2) 𝑑𝑖 = ||𝐀𝑥𝑖 − 𝑦𝑖||.Solving for 𝐀 that minimizes the distance 𝑑𝑖 between the two shape models, as shown in [12], will compute the parameters of the required projective transformation. This process is repeated on the transformed shape until the change in the transformed points is not significant. Fig. 4 shows the registration results for samples from the four builds at different layers.Figure 3. Laser beam power setting in watts as a function of layer area for the four builds.Figure 4. Examples of peak temperature during the scan of different layers. The first row is for the first build as described in figure 3.

Examples of peak temperature during the scan of different layers. The first row is for the first build as described in figure 3. The effect of changing laser power is very clear in the second row when we reduced the laser power of the small cross-sectional area too much. The last row shows a uniform peak temperature.Analysis of ResultsThe system setup is shown in Fig. 2. The system consists of the new “open architecture” LS machine, thermal camera and a 2 in flat gold mirror as described in the data collection section above. Four builds were conducted in the new “open architecture” LS machine. Builds are done using PA 650 nylon 12 laser sintering material [13]. Builds are done in the vertical orintation as shown in Fig. 3. The first build is conducted with fixed laser beam power, while the other three builds are conducted with variable laser beam power as a function of layer cross-sectional area as shown in Fig. 3.

3. The first build is conducted with fixed laser beam power, while the other three builds are conducted with variable laser beam power as a function of layer cross-sectional area as shown in Fig. 3. Four different videos are recorded during the build of the test parts with four different scan parameters settings.Thermal data for the builds are segmented using template to image registration method as explained before. Samples of segmented images for every build are shown in Fig. 4. The average peak temperature in every layer is used as a measure. Fig. 5 shows the plots of the average peak temperature versus layer number for the four builds. These plots show the effect of changing the laser power on the average peak temperature. The first plot, for the build with fixed laser power, shows that layers with small area have higher temperature values and the very large area has the lowest temperature. With changing the laser power, as detaild in Fig.

With changing the laser power, as detaild in Fig. 3, we can see how this affect the average peak temperarure in the second, third, and last plot. The results indeed confirmed that we can get a uniform peak temperature as shown in the last plot. This observation has been confirmed by Fig. 6 that shows that the very small area has the highest temperature for the build without control, but with controlling the laser power we get a constant peak teperature for different areas. These results approved that the layer area can be used as a control signal to control the laser beam power to maintain constant peak temperature throughout the part. This produced more consistent part properties.DiscussionThis paper addresses the problem of controlling the 3D printing process. We presented an approach for polymer laser sintering process control based on part geometry. The proposed control system utilized information from layerwise geometry models for the part to control the fabrication process.

The proposed control system utilized information from layerwise geometry models for the part to control the fabrication process. We used a single layer geometry measure, which is the layer scanned area, to pre adjust the laser beam power for every layer. We empirically adjusted the laser power based on the layer cross-sectional area. Results from four different builds show that we can get a uniform average peak temperature throughout the part. For future work we will use the collected data to estimate a closed form relation between the layer area, the peak temperature, and the laser beam power to precisely control the laser power. Also, we will study the effect of the pre-scan powder bed temperature on the fabrication process.Figure 5. Average peak temperature in every layer depends on the cross-sectional area and the laser beam power. The lower plot for the last build approved that we can get a uniform peak temperature by controlling the laser power.498Figure 6.

F at Ford MJtor Crnp;':mySean B. O'Reilly Ford :fXk)tor Carpany Alpha Simultaneous Engineering 24500 Glendale Avenue Redford, Michigan 48239Intrcrluctian. Ford's effort in Free Fonn Fabrication (FFF) began in 1987with the fontation of an internal consortium, Calp)sed of a dozen differentactivities, whose purpose was to evaluate and apply, where possible,thiserrerging tecl:mology. Each of the consortium members agreed to contributesare rroney and,rrore irrportantly, one or rrore people to work on thisproject. By the following year (1988)the first rrachine, an SLA-1, had beeninstalled at the Alpha Msmufacturing Develq:ment Center and fomed thecornerstone of the FFF Lab.An Early Success. By the end of the year a number of successful parts hadbeen built. One of the rrost drarratic was a rocker ann for the 1992 5Lengine. The story of this ann bas been told several t:i.Ires in articles andpresentations but, nonetheless, bears repeating.

One of the rrost drarratic was a rocker ann for the 1992 5Lengine. The story of this ann bas been told several t:i.Ires in articles andpresentations but, nonetheless, bears repeating. A yaung engineer haddesigned this ann and had sent the design out to suppliers for quotes.(Thisparticular part is purchased fran outside.) Nonral practice at that t:i.Ire wasto use a 2-D detailed drawing to describe the part to be quoted. The initialround of quotes had already been carpleted :based on the infontation containedin the drawing. Four suppliers had been contacted and asked to quote. Threeof the four had responded with per-unit prices and the fourth had care backwith a "No-Bid". The engineer was assigned to the FFF lab, built a 3-D Solidrrodel of the part and then used the rrodel to build a stereolithography168prototype. It was only upon inspection of the prototype that he discovered aproblem with the design.

It was only upon inspection of the prototype that he discovered aproblem with the design. He corrected the design and built a second prototypeto verify that the problem had been solved. Since the design had changed, asecond round of quoting was initiated. This time, however, plasticstereolithography prototypes were sent to the suppliers in place of reviseddrawings. Each of the suppliers carre back with a lower quote than on thefirst round and the supplier who had IIno-bid ll first time around, carre in ata price roughly 50 cents lower than anyone else. This reduction amounted toseveral million dollars of savings per year.Fundan:ental Value of FFF. The rocker arm story illustrates the basicvalue of FFF technology. The ability to take a ccnputer rrodel and build anaccurate prototype of it in a few hours or days provides a new, highbandwidth medium of camunication for infoIltEltion about parts.

The ability to take a ccnputer rrodel and build anaccurate prototype of it in a few hours or days provides a new, highbandwidth medium of camunication for infoIltEltion about parts. Consider theprocess by which 2-D drawings are produced today:the engineer conceptualizesthe part in his head; rrekes a few rough 2-D sketches and passes these alongto a designer or detailer who rrust first get an accurate concept of thepart,then, conceptually pass planes through the conceptualized part atvarious angles and then accurately conceptualize how these cross-sectionswould look;then he begins drawing these conceptualized cross-sections. Thisis a very carplex and error-prone process. once the drawings are carplete,they are sent to a nanufacturer who rrust then reverse the process conceptualI yand reconstruct the original design intent fran the drawn cross-sectionalinfoIltEltion - an even rcore carplex and error-prone process.

All of thisconceptualis carpletely short -circuited by FFF. The FFF prototypebecanes the prirrary rrethod of carmunicating all of the geanetric intricacies169of the part.In a sense, representations of the part are replaced by thepart itself. It is never easy to quantify the benefits of inpravedcarrnunications, but the millions in savings spawned by the plastic rockerann certainly indicate that 11 •• there I s gold in them thar hills! ! ! ... 11FFF and P.r:ototyping. Prototypes have been in the mainstream of invention,engineering and product developrent since Alley Oop first laid chisel tostone in search of the wheel.In the PaSt,the role of prototyping has beenseverely limited by our inability to produce prototypes rapidly and costeffectively. FFF holds out the ProsPect of being able to provide prototypesat virtually every stage of product developrent. Let Ire en:phasize that thisability is not here today, but is caning.

FFF holds out the ProsPect of being able to provide prototypesat virtually every stage of product developrent. Let Ire en:phasize that thisability is not here today, but is caning. Prototypes seem to fall into threemain categories, related to the stage of developrent of the product:I1touchfeel" prototypes,I1fonn, fit,limited function l1 prototypes and fullyfunctional prototypes.I1Tauch- Feel 11 Prototypes: Like the rocker ann cited above,these prototypesare used in the early stages of design for a IlUll'ber of purposes - conceptverification, carrnunication of design intent, even for verifying the CADrepresentation of the part. We have seen numerous examples where the FFFprototype showed up a flaw in the CAD m:xlel that was not detectable on thescreen.

We have seen numerous examples where the FFFprototype showed up a flaw in the CAD m:xlel that was not detectable on thescreen. This is the area where FFF currently has its greatest i.rrpa.ct.11 Fonn, Fit, Limited Function l1 Prototypes: Canponents designed by differentactivities nust often be assembled to fonn an integrated system. The abilityto build each carponent quickly and accurately is of trerrendous value inassembly studies,in finding interferences and in discovering problems earlyenough that they are easily handled. Again FFF prototypes are having a170trerrendous in:pact in this arena. Fig. 1. shows a partial engine block, madein an SLA 500.'Ibis block was sent to the plant which was to manufacture it.The plastic block was used to check aut tooling, fixturing, rraterial handlingequipre:nt and dunnage. A number of potentially costly problems were discoveredand corrected before the final specification for the plant equipre:nt. Fig.2.

A number of potentially costly problems were discoveredand corrected before the final specification for the plant equipre:nt. Fig.2. shows an experirrental intake manifold rrade on a Cubital solider. 'Ibismanifold was attached to an engine and nm through several power cycles todetennine the effect of the design on engine perfonrance. We have made manysuch manifolds on Cubital and SLA rrachines. The ability to actually test anew design within a few days is a capability that our engineers have neverhad before.'Ibis is an exarrple of a "limited functionality" prototype.Several of our divisions now make this type of prototype where thefunctionality is prirrarily due to the gearetry. Mmifolds,throttle bodies,pump housings, alternator housings and AC housings are all exarrples of thiskind of application."Full Functionality" Prototypes: For the rrost part, FFF-generated partsdo not have the rraterial properties required for full functionality.

In thiscase,the FFF part is used as a m:::del or patte:r:n in a more traditional typeof manufacturing process. Fig. 3a shows a stereolithography transmissioncase and Fig. 3b shows the aluminum casting that was made in our experirrentalfoundry. Figs. 4a and 4b show a stereolithography crankshaft and the iron onrrade fran it. In both instances, our engineers have gotten functionalprototypes in a fraction of the tine it would have taken using conventionalnEthods. Sand, plaster and investnEnt casting of nEtals and molding of saneplastics (e.g. silicone rubber molding of polyurethane) are all sped up171dra:natically by the use of FFF models as rattems. Significant time reductions(at least 50%)in the prototyping process are a.1Irost universal.FFF and Tcx:iling. A recent study has estimated that there are 1200 PARTSTHAT mST BE DESIGNED AND :MANOFACIURED FOR EACH PART THAT GOES INTO THEAVERAGE FORD VEHICLE.

A recent study has estimated that there are 1200 PARTSTHAT mST BE DESIGNED AND :MANOFACIURED FOR EACH PART THAT GOES INTO THEAVERAGE FORD VEHICLE. Tooling Ifixtures I ma.terial bandling and durrnage allfall into this catego:ry. A similar study by GM came up with an estimate ofabout 1800 such r:arts. Since work on these r:arts can't seriously begin untilthe product design is reasonably stable, it seems evident that the applicationof FFF to this arena will be essential to reduce time to ma.rket. We arecurrently carrying aut work to develop ways of ma.king tools (molds and dies)directly fran FFF r:arts. Fig. 5 shows an FFF r:art in the right foreground.It is actually a sand core which will be assembled into a sand mold to ma.kean aluminum engine block. This r:articular core will be an internal oil drainPassage. To the left of the r:art is a stereolithography prototype of ther:art sitting in a mold half.

This r:articular core will be an internal oil drainPassage. To the left of the r:art is a stereolithography prototype of ther:art sitting in a mold half. This was produced by designing the mold half ona CAD station with the r:art enibedded. In the background are two epoxy moldhalves which were famed as follows:the right mold half was famed directlyfran the stereolithography piece. The left half was then famed fran theright half with the r:art inserted.FFF Limits. These few exan:ples illustrate the variety of r:arts andapplications that we have carpleted. Where applicable FFF gives a trerErJdousadvantage in te:r::ms of time savings and cost savings. However, not all of ther:arts the Ford JXbtor Carpany ma.kes can be successfully prototyped tooayusing FFF. The most obvious limit is the size of a r:art that can be built inone piece.:M3ny of our r:arts rrust be built in two or more pieces and then172joined.

The most obvious limit is the size of a r:art that can be built inone piece.:M3ny of our r:arts rrust be built in two or more pieces and then172joined. In rrany cases this is a perfectly satisfactory process, prcxiucingacceptable prototypes. However,this approach doesn't work, for example,the area of large sheet rretal parts. No machine available tc:x:lay can build adurable hocxi or fender aut of any material. Along with size,the layeringprocess appears to be an i.rrped.i.rrent in very thin but extended parts.Coopared to rrore traditional processes for making prototypes, FFF is nothingshort of a rni:r:acle.

After a while however, speaking of limits, one begins tobecare impatient with the build speed of tc:x:laY's FFF machines.(An exampleof the infinite perversity of ham sapiens? - They're NEVER satisfied.)Given the :r:apid (and seemingly accele:r:ating) pace of develq:roent in theFFF arena and the fact that the industry is still in its infancy, we canexpect to see rrany of tc:x:lay's limits avercare in very short order.inssion Case1MCastTransmis ionusIron Crankshaftspart: FFF model mod(:;l of mold made.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFiber Encapsulation Additive Manufacturing: Materials for Electrical Junction FabricationB. Xia, M. Saari, B. Cox, E. Richer, P. S. Krueger, and A. L. Cohen*Laboratory for Additive Manufacturing, Robotics, and Automation (LAMRA), Department of Mechanical Engineering, Bobby B. Lyle School of Engineering, Southern Methodist University, Dallas, TX 75275Corresponding authorAbstractFiber Encapsulation Additive Manufacturing (FEAM) is a novel 3-D printing process that permits the printing of electromechanical and electronic devices within a single, affordable machine. A key challenge of FEAM is creating robust and reliable electrical junctions between encapsulated wires, enabling more complex devices and circuits to be fabricated.

A key challenge of FEAM is creating robust and reliable electrical junctions between encapsulated wires, enabling more complex devices and circuits to be fabricated. We present current efforts to explore and characterize several different methods for creating junctions: solder, solder paste, and a custom-formulated electrically conductive polymer composite. All three methods are analyzed in terms of printability, material compatibility, repeatability, and performance.IntroductionIn recent years, increasing research attention has been focused in the area of additive manufacturing (AM) of electronics and electromechanical devices [1-3]. These investigations attempt to utilize the intrinsic ability of AM to rapidly fabricate complex customized geometries, enabling designers to pursue sophisticated designs that were previously impossible.

Additionally, such extensions of AM promise to achieve a higher level of functionality and performance, significant reductions in size, weight, and cost when compared to traditional manufacturing techniques.A rapidly-growing field in robotics is soft robotics. Soft robots have several important advantages over conventional rigid robots [4]. Made from soft materials, such robots can function safely in close proximity to people, which is an advantage in co-robot systems, allowing them to work with people, and aid them in a multitude of tasks. Also soft robots can grasp and manipulate delicate irregular objects in ways that rigid robots cannot. For example, a soft robotic gripper can catch an uncooked egg without damaging the shell [5]. Additionally, soft robot parts can provide more degrees of freedom, enable more complex motion, and undergo shape transformations. These attributes can allow them to work in limited spaces, such as search and rescue operations after earthquakes.

These attributes can allow them to work in limited spaces, such as search and rescue operations after earthquakes. Soft robots typically have many degrees of freedom, which while enabling a large range of shapes and motions, may necessitate a large number of distributed actuators. Moreover, they may incorporate distributed sensors, for example, for proprioceptive feedback and in a tactile sensing “skin”.1345Motivated by the benefits of 3-D printed electronic and electromechanical devices and by the needs of soft robots, Fiber Encapsulation Additive Manufacturing (FEAM) has been developed by researchers at the Laboratory for Additive Manufacturing, Robotics, and Automation (LAMRA) at Southern Methodist University [6]. FEAM is a novel additive manufacturing process in which a fiber and a matrix are co-deposited simultaneously within a single printer along straight or curved paths.

FEAM is a novel additive manufacturing process in which a fiber and a matrix are co-deposited simultaneously within a single printer along straight or curved paths. In FEAM the fiber can be a robust, high-conductivity metal wire (assumed here) for electronic/electromechanical applications, while the matrix can be a soft material, e.g., for soft robotics and wearable electronic applications, or a hard material, as required.To develop FEAM, three key capabilities are required: wire encapsulation and centering, starting and stopping the wire, and forming electrical junctions between wires. Significant progress has been made and already described with regard to the first two capabilities [7]. Based on earlier work, a number of devices have been produced including a loudspeaker, inductive sensors, a solenoid actuator, membrane switches, and a capacitive force sensor, as shown in Figure 1.

Based on earlier work, a number of devices have been produced including a loudspeaker, inductive sensors, a solenoid actuator, membrane switches, and a capacitive force sensor, as shown in Figure 1. This paper will focus on printing electrical junctions, which enable far more complex circuity and devices.Figure 1. Devices printed using Fiber Encapsulation Additive Manufacturing.1346Printing Electrical JunctionsIndividual parts such as the simple coil shown in Figure 1 (top middle) or even more sophisticated coils (top right) may be printed with a single continuous length of wire. However, in context of printing all but simple circuits, multiple discrete straight or curved segments of wires will be required. Connecting discrete segments of wires in the same layer or across multiple layers thus becomes important to the FEAM process.

Connecting discrete segments of wires in the same layer or across multiple layers thus becomes important to the FEAM process. Additionally, in order to automatically, monolithically print an entire pre-assembled system which may include electromechanical devices, actuators, sensors, interconnects, processors, and power sources, junctions between printed wires are required.However, it is not trivial to engineer junctions to work within a FEAM system. To reliably fabricate an electronic or electromechanical device, junctions need to have low resistivity and high current capacity so electrical components can operate with good performance and acceptable efficiency. It is also important that the junction be robust and durable, especially if complex motion or shape transformation of the device is part of the design objective. The material to fabricate the junctions needs to have good printability and compatibility with the surrounding material.

The material to fabricate the junctions needs to have good printability and compatibility with the surrounding material. Repeatability must also be considered to guarantee the quality of the fabricated system. Moreover, other aspects such as adhesion of the junction material to wire, cost, and printing time need to be included in the choices for electrical junction fabrication.The wires or the wire and electrical elements to be connected may lie in the same layer or in adjacent layers. In an effort to model these kinds of printing conditions, junctions of two types of structures are tested and discussed in this paper. An intra-layer junction is one in which two or more wires lie in the same layer as shown in Figure 2 (a). An inter-layer junction is where wires lie in neighboring layers shown as Figure 2 (b).(a) Intra-layer junction (b) Inter-layer junctionFigure 2.

An inter-layer junction is where wires lie in neighboring layers shown as Figure 2 (b).(a) Intra-layer junction (b) Inter-layer junctionFigure 2. Schematic junctions.1347To satisfy the performance requirements for junctions, various materials have been investigated. Three following candidate materials and their printing methods are explored: electrically conductive polymer composite (ECPC), solder, and solder paste.Electrically Conductive Polymer CompositeA variety of custom ECPC materials have been explored by our group [8]. They consist of an electrically conductive metal powder and a polymer matrix. Polymers in ECPC are intended to bind metal powder and enable material extrusion. The metal powder chosen was silver-coated nickel. Three types of polymers were tested for ECPC formulation including Kraton D1161 P (Kraton Performance Polymers Inc.

The metal powder chosen was silver-coated nickel. Three types of polymers were tested for ECPC formulation including Kraton D1161 P (Kraton Performance Polymers Inc. Houston, TX), Crystalbond 555 (Aremco Products Inc., Valley Cottage, NY), and beeswax.Among the various materials explored, beeswax was unable to bind to the metal powder sufficiently well in order to achieve sufficient blending, as seen in Figure 3. The ECPC it produces was very fragile, and did meet the mechanical strength requirements for the FEAM process.Figure 3. ECPC made of beeswax and silver coated nickel powder.Crystalbond 555 was more readily blended with the metal powder at the same volume fraction as the Kraton D1161 P. The resulting ECPC was fabricated quickly (within five minutes) by blending the components at elevated temperature, even when the volume fraction of metal powder approached 35%.

The resulting ECPC was fabricated quickly (within five minutes) by blending the components at elevated temperature, even when the volume fraction of metal powder approached 35%. In comparison, ECPC produced with Kraton D1161 P, a thermoplastic elastomer, could only be mixed to a volume fraction of metal powder as large as 29% while simultaneously maintaining printability. Additionally, Kraton D1161 P ECPC took around two hours to formulate the same volume as that made with Crystalbond 555. Crystalbond 555 however had some difficulty binding to the metal powder. The resulting resistivity of the ECPC made from this kind of polymer was almost ten times higher than the ECPC made from Kraton D1161 P with the same volume fraction of metal powder, as shown in Table 1. The ECPC made of Crystalbond 555 was less readily extruded than that made of Kraton D1161 P as well.

The ECPC made of Crystalbond 555 was less readily extruded than that made of Kraton D1161 P as well. Comparing the two types of material,1348it can be concluded that the ECPC made of Kraton D1161 P has better abilities to fabricate ECPC- based junctions in the FEAM system than that made of Crystalbond 555. Thus, only Kraton D1161 P is investigated further here.Volume Fraction23%25%27%29%30%33%35%Crystalbond 555 (Ω-cm)nonconductive 0.571 0.568 0.560 0.558 0.551 0.547Kraton D1161 P (Ω-cm)0.1480.038 0.026 0.019cannot be fabricatedTable 1. Resistivity of ECPC made from TPE D1161P and Crystalbond 555 vs. volume fraction of metal powder.Since metal powder is providing the electrical conductivity in ECPC, it is expected that increasing the volume fraction of metal powder will result in the resistivity decreasing somewhat. This was observed experimentally, as shown in Table 1 and Figure 4.

This was observed experimentally, as shown in Table 1 and Figure 4. The reason for the rapid decrease at the beginning and the slower decrease of resistivity at higher powder volume fraction is supported by the fact that once electrical percolation has been achieved, relatively small decreases in resistivity are expected as more powder is added. When the volume fraction of metal powder is large enough, approximately 25%, the powder in ECPC has sufficient metal particulate to permit percolation allowing current to flow. After that point, the slower decreasing rate of resistivity is limited by the binder capacity and the resistivity of the metal powder.Resistivity vs Volume Fraction of Metal PowderC P C E f o y t i v i t s i s e R)m c - Ω (0.16 0.14 0.12 0.1 0.08 0.06 0.04 0.02 022%26% Volume Fraction of Metal Powder24%28%30%Figure 4.

Resistivity vs Volume Fraction of Metal Powder.Concurrently, since printability (via material extrusion) is determined in part by material rheology, it is expected that printability will behave in the opposite fashion to conductivity with respect to increasing volume fraction of metal powder. Indeed, it was found that printability decreases dramatically as volume fraction of metal powder increases. When the volume fraction becomes 30%, the ECPC formulation has a rough surface texture, is friable, and can no longer be printed through a small nozzle, as shown in Figure 5. So by balancing these two opposing properties, the1349optimal volume fraction of silver-coated nickel powder in Kraton D1161 P ECPC used to print electrical junctions was approximately 29%.(a) 29% metal powder. (b) 30% metal powder.Figure 5.

(b) 30% metal powder.Figure 5. ECPC filament made of Kraton D1161P with 29% and 30% metal powder.Under constant tensile load applied to ECPC filament, it was observed that resistivity decreased with time, as shown in Figure 6. When a constant load is applied to the filament, strain will be generated. The strain will increase rapidly at first and slow down when approaching the strain relaxation limit. During increased strain relaxation, resistivity decreased. We hypothesize that as the filament elongates, the metal powder particles in the ECPC filament become pressed more closely together, increasing the number and area of conductive paths between particles. However, it was not possible to determine relationship between strain and resistivity from the present resultsResistivity vs. Time at Constant Tensile Load (20g))m c - Ω( y t i v i t s i s e R0.1 0.09 0.08 0.07 0.06 0.05 0.04 0.03 0.0205001000 Time (seconds)15002000Figure 6.

Time at Constant Tensile Load (20g))m c - Ω( y t i v i t s i s e R0.1 0.09 0.08 0.07 0.06 0.05 0.04 0.03 0.0205001000 Time (seconds)15002000Figure 6. Time dependent effects of constant loading on resistivity of an ECPC filament comprising Kraton D1161.In addition, the experimental results show that intra-layer and inter-layer ECPC-based junctions require a very short time to be printed. 3 seconds is enough time to deposit an inter-layer or intra- layer junction; however, printing time varies with the geometry of the parts. The typical resistances of these two kinds of junctions are 250 mΩ and 110 mΩ, respectively. The difference in resistance is attributed partly to the distance between wires (~250 µm center-center for inter-layer junctions, vs. ~450 µm for intra-layer junctions) while the relative high resistance of the junctions is likely caused by poor wire adhesion to ECPC and/or voids in the ECPC.1350(a) Intra-layer junction.

~450 µm for intra-layer junctions) while the relative high resistance of the junctions is likely caused by poor wire adhesion to ECPC and/or voids in the ECPC.1350(a) Intra-layer junction. (b) Inter-layer junction.Figure. 7. ECPC junctions.The Kraton D1161 P-based ECPC was printed using the Thermoplastic Elastomer Additive Manufacturing (TEAM) system developed at LAMRA [8]. The TEAM system consists of a constant pitch miniaturized screw extruder, inspired by industrial screw extruders, mounted as a material extrusion extruder on a 3-D printer. Other research [9] has shown this system able to 3-D print a wide variety of TPEs including Kraton D1161 P.The TEAM extruder design is shown in Figure 8. ECPC pellets were fed into the extruder through a pellet hopper, and are transported through the barrel to the extruder by the screw rotated by a geared stepper motor.

ECPC pellets were fed into the extruder through a pellet hopper, and are transported through the barrel to the extruder by the screw rotated by a geared stepper motor. The hot-end block and barrel are heated by a cartridge heater and band heater, combined with a heat sink to produce the desired temperature gradient to drive flow. The ECPC was extruded out of the nozzle into prefabricated junction cavities (in which the wires were embedded in non-conductive material printed with FEAM), producing conductive junctions.Figure 8. TEAM miniature screw extruder schematic.1351On a volumetric basis, the materials used in ECPC junctions have the lowest price among all three type of junctions compared in this paper, but other junctions may use less material. The material melting point of ECPC is relatively low, and ECPC junctions can be printed with reasonable repeatability if the junction is well designed.

The material melting point of ECPC is relatively low, and ECPC junctions can be printed with reasonable repeatability if the junction is well designed. The relatively short time required to print an ECPC junction makes for printing an entire electronic/electromechanical system. Thus even though ECPC junctions are considerably less conductive than others we have investigated, they remain an attractive option.the process very efficient when utilizedSolder-based JunctionsSolder is a very common material used to fabricate electrical junctions. Several experiments have been conducted by multiple groups, such as printing solder with material extrusion systems [10], and directly printing solder to build circuit board surface traces [11].

Several experiments have been conducted by multiple groups, such as printing solder with material extrusion systems [10], and directly printing solder to build circuit board surface traces [11]. It is clear that, at low cost, solder can fabricate electrical junctions that have very low resistance, and are robust and have good mechanical strength.However, other aspects such as printability and material compatibility need to be investigated to decide whether solder is a good fit for fabricating electrical junctions in combination with FEAM. To test solder junctions, two printing methods were investigated: molten solder extrusion and forced convection melting.Molten solder extrusion is a method explored for printing junctions with a syringe extruder as seen in Figure 9. The solder is put into the syringe, a wraparound heater heats the solder to above its melting point, and then the solder is injected through a heated needle onto the wires to form the junction.Figure 9.

Molten solder extrusion. Figure 10. Forced convection melting.Forced convection melting is a similar method; but in this case, hot air is convected to the solder by a ZT-2-MIL hot air pencil (Zephyrtronics, Pomona, CA) as shown in Figure 10. After the hot air pencil is preheated to the solder’s melting point, hot air is aimed at solid wire solder beneath it. The melted solder detaches and drops onto the wires to form the junction.1352In52/Sn4858Bi/42Sn97In/3AgMelting point (°C)118138143Minimum solder ball dia obtained (mm)0.830.990.93Resistance of fabricated junctions (mΩ)6.97.16.8Table 2. Solder used to fabricate electrical junctions.For each of these printing methods, three types of solder were tested, and the resulting data is listed in Table 2.

Solder used to fabricate electrical junctions.For each of these printing methods, three types of solder were tested, and the resulting data is listed in Table 2. From experiments, it can be concluded that because of the relatively large surface tension of the melted solder, achieving a solder ball small enough to avoid damage to the surrounding (thermoplastic) polymer is difficult.In forced convection melting, the accuracy of the final solder position is very poor due to the bulk motion of the hot air dispersing the solder. Thus, even though solder-based junctions have very low resistance, are robust and durable, print quickly and are cost-effective, such junctions do not easily meet the requirements to fabricate electrical junctions as part of a FEAM system.Solder Paste Deposition and Reflow JunctionsCompared with solder, solder paste is another common material used to fabricate electrical junctions.

As with solder, low electrical resistance, robust and durable mechanical strength can be achieved. However, solder paste consists of very small solder balls immersed in a liquid solder flux and needs to be dispensed. A pneumatic dispenser can be used to controllably dispense solder paste onto wires as the first step in forming a junction.To control the amount of the solder paste used for fabricating a junction, a dispenser controller (DS-982, Taiwan Tech & Material, Taiwan) was utilized. The amount of solder paste injected out of the syringe is a function of the air pressure, needle size, and dispensing time. Since needle size affects the pressure gradient in the solder paste required to drive the flow, and the position of the solder paste relative to the wires, this is the first parameter considered. Needles with IDs from 0.2 mm to 0.8 mm were tested.

Needles with IDs from 0.2 mm to 0.8 mm were tested. When needle size increased, dispensing time decreased as expected, but the accuracy of the dispensed solder paste’s position decreased. If the needle was too large, solder paste would flow under hydrostatic forces. Conversely, if the needle was too small, solder paste would not be extruded with the available pressure.1353To balance the effects of dispensing speed and the resulting positional accuracy, a needle of 0.33 mm ID was chosen to dispense the solder paste. To match the needle size, air pressure was set at 65 psi and the dispensing time was set to 0.5 seconds. The result was a roughly spherical solder paste ball of around 0.45 mm diameter deposited onto the wires. Using this dispensing method, three reflow methods were tested independently: soldering iron, hot air pencil, and laser. Soldering irons are common and inexpensive tools used for soldering.

Using this dispensing method, three reflow methods were tested independently: soldering iron, hot air pencil, and laser. Soldering irons are common and inexpensive tools used for soldering. Since the reflowing apparatus must be integrated into the FEAM printer, and it may need to form a large number of junctions within a single device without stopping, it is of vital importance to keep the soldering tips clean after contact with the solder paste. To study this requirement, an assortment soldering tips made of PTFE and polished aluminum, and having different shapes and sizes as shown in Figure 11 were tested.Figure 11. Custom-made soldering tips.Good results producing intra-layer junctions were attained when heating the solder paste for two seconds, however, the tips still damaged the surrounding material.

Custom-made soldering tips.Good results producing intra-layer junctions were attained when heating the solder paste for two seconds, however, the tips still damaged the surrounding material. Additionally, inter-layer junctions could not be fabricated in a straightforward manner with this method: nearly 50% of inter-layer junctions could not be formed during experiments, and in the case of successful inter- layer junctions (those having low resistance and good mechanical integrity), the surrounding material was irreparably damaged. Therefore, this approach was not considered practical.Figure 12. Solder paste-based junction reflowed by soldering iron.Hot air sources may be used to reflow solder paste. To evaluate this approach a hot air pencil reflow system was constructed, comprising a computer-controlled X/Y positioning stage as shown in Figure 13.Figure 13.

To evaluate this approach a hot air pencil reflow system was constructed, comprising a computer-controlled X/Y positioning stage as shown in Figure 13.Figure 13. Hot air reflow system.1354With this method, the temperature of the hot air was set to the melting point of the solder paste. The main parameter varied was the size of the air pencil nozzle. If the nozzle was too small, the mean flow velocity would be too large, and the solder paste dispensed on the wires could be easily displaced by the air jet. If the nozzle was too large, the surrounding 3-D printed material would be damaged by the hot air because it had a lower melting point than the solder paste. To balance the two effects, nozzles with an ID range from 1 mm to 5 mm were tested. The best results were obtained with a 2-mm nozzle and a reflow time of 5 seconds, as shown in Figure 14. Even though the surrounding material is barely damaged, there is some residue inside the cavity.

Even though the surrounding material is barely damaged, there is some residue inside the cavity. Nonetheless, the results show that it can meet the requirements to fabricate electrical junctions as part of FEAM.(a) Intra-layer junction. (b) Inter-layer junction.Figure 14. Solder paste-based junction reflowed by hot air (intra-layer and inter-layer).Laser induced thermal reflow is a less common method to reflow solder paste. The laser reflow system constructed for the investigation, shown in Figure 15, uses a computer-controlled X/Y positioning stage to control the position of the laser focal point. The laser system used included a multimode pigtail diode laser (Innolume GmbH, Dortmund, Germany), with a laser driver and mount (Arroyo Instruments LLC, San Luis Obispo, CA), and fiber collimator and lens (Thorlabs, Inc., Newton, NJ).Figure 15. Automated laser solder reflow system.The solder paste dispensed on the wires is subsequently heated by the focused laser beam.

Automated laser solder reflow system.The solder paste dispensed on the wires is subsequently heated by the focused laser beam. Heating time and laser current are the two primary variables which affect the quality of the junction. When the current was made too large, heating was too intense and rapid and caused spattering of the solder. If it was made too small, the heating time was extended. A current setting of 5 A with heating time of 2 seconds resulted in the junctions of the best quality for both intra-layer and inter- layer junctions, as shown in Figure 16. There is no observed difference in junction quality between inter-layer and intra-layer junctions when printing solder paste-based, laser-reflowed junctions, as the solder paste will directly form a solder ball around the two wires.

However, even though the1355processing time is only two seconds for the laser reflow, dispensing solder paste and subsequently moving the junction under the laser significantly can increase the total printing time parts with many junctions.(a) Intra-layer junction. (b) Inter-layer junction.Figure 16. Solder paste-based junction reflowed by laser (intra-layer and inter-layer).After comparing all three methods of solder paste reflow methods, laser reflow has overall the best characteristics for obtaining electrical junctions in use with FEAM. Even though minor residue inside the cavity remains and device costs are relatively high, laser reflow still outperforms the other methods tested to date. Hot air reflow can also produce acceptable junctions for FEAM. Lastly, soldering iron reflow does not appear to be viable.Three types of solder paste have been tested and all produce acceptable electrical junctions using the methods above.

Lastly, soldering iron reflow does not appear to be viable.Three types of solder paste have been tested and all produce acceptable electrical junctions using the methods above. However, since laser reflow can produce junctions of the best quality as discussed above, the properties of several solder pastes were compared by using this reflowing method, as shown in Table 3.

However, since laser reflow can produce junctions of the best quality as discussed above, the properties of several solder pastes were compared by using this reflowing method, as shown in Table 3. As shown, overall 96.5Sn/3Ag/0.5Cu (Version B) is the preferred solder paste based on experimental results.58Bi/42Sn96.5Sn/3Ag/0.5Cu (Version A)96.5Sn/3Ag/0.5Cu(Version B)Melting point (°C)138217-219217-219Junction resistance (mΩ)7.317.267.18Cost (USD/cm3)1.252.581.75Post-reflow cleaningNoneNoneNoneJunction mechanical strengthGoodExcellentExcellentTable 3: Comparison of junctions made with three types of solder paste and laser reflow.1356Comparison of the Three Materials and SummaryFor all three types of materials discussed and their corresponding printing methods, a detailed comparison is shown in Table 4.

Accordingly, it can be concluded that solder is not a good material to fabricate electrical junctions in FEAM due to its poor printability, repeatability, and process compatibility. ECPC and solder paste satisfy all the performance requirements for printing electrical junctions. Overall, solder paste may be described as the best material for fabricating electrical junctions. However, ECPC should still be considered as it can meet the requirements and has certain advantages such as faster processing.Table 4. Comparison of properties of ECPC, solder and solder paste.AcknowledgementsThis paper is based on work supported by the National Science Foundation under grant no. 1317961. Research was performed at the Laboratory for Additive Manufacturing, Robotics, and Automation, Department of Mechanical Engineering, Southern Methodist University, Dallas, Texas.

1317961. Research was performed at the Laboratory for Additive Manufacturing, Robotics, and Automation, Department of Mechanical Engineering, Southern Methodist University, Dallas, Texas. The authors appreciate for their assistance Todd Danner, Daniel Porter, Samantha Reeve, and Jenya Pryadkin.Author Disclosure StatementThe authors have no financial conflicts.1357References[1] Espalin, David, et al. "3D Printing multifunctionality: structures with electronics." The International Journal of Advanced Manufacturing Technology 72.5-8 (2014): 963-978. [2] Ko, Seung H., et al. "All-inkjet-printed flexible electronics fabrication on a polymer substrate by low-temperature high-resolution selective laser sintering of metal nanoparticles." Nanotechnology 18.34 (2007): 345202.[3] Muth, Joseph T., et al. "Embedded 3D printing of strain sensors within highly stretchable elastomers." Advanced Materials 26.36 (2014): 6307-6312.[4] Majidi, Carmel.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFiber traction printing--a novel additive manufacturing process ofcontinuous fiber reinforced metal matrix compositeXinWang, Xiaoyong Tian, Dichen LiState Key Laboratory of Manufacturing Systems Engineering, Xi’an Jiaotong University, ChinaAbstractMetal matrix composites (MMCs) are materials which have been widely used in the aerospace andautomobile industries since the 1980s and have been classified as hard-to machine materials. This manuscriptproposes a novel additive manufacturing process of continuous fiber reinforced metal matrix composite -- fibertraction printing (FTP). The composites with complex structures can be directly manufactured via FTP whichutilizes the wetting force and capillarity force to control the flow of melting matrix.

The composites with complex structures can be directly manufactured via FTP whichutilizes the wetting force and capillarity force to control the flow of melting matrix. The craft is proceedingwithout extra pressure device and atmosphere protection device and substantially decreases the cost.

The craft is proceedingwithout extra pressure device and atmosphere protection device and substantially decreases the cost. Thismanuscript introduces the proof-of-concept prototype and the ability to control the flow of melting matrix andfiber distribution through this process yields a flexible manufacturing route to fabricating 3D metal matrixcomposite parts with full density and complex geometries.Keywords: metal matrix composite; 3D printing; additive manufacturing;IntroductionManufacturing technologies in extreme environments have drawn more and more scientists' attentionespecially for space exploration.1 In particular, 3D printing techniques have potential to be deployed in the spaceenvironment to enhance safety levels by providing replaceable parts and decrease launch mass.2-7 Hence, metalsor materials with higher performance are required.

Nevertheless, the conventional metal based 3D printingmethods, such as SLS, EBM and SLM,8 are challenging to be operated in space7-9 and shrinking and balling ofmolten droplets occur in microgravity due to the large surface tension.9,10 Thus, a 3D printing method utilizingfiber pulling process is proposed in this paper in order to prepare metal matrix composites (MMC). Fiber TractionPrinting (FTP) is an additive manufacturing method taking advantage of wetting behavior between fibers andmetal matrix to pull the molten metal out of printhead and then deposit the composite filaments layers uponlayers. The wetting forces are used to overcome the surface tension of metal, thus molten metal canhomogeneously enfold the fiber and spread along fiber direction.

The wetting forces are used to overcome the surface tension of metal, thus molten metal canhomogeneously enfold the fiber and spread along fiber direction. To demonstrate the fundamental mechanism ofthis approach, we will show a proof-of-concept FTP prototype of a simple and low temperature C f/Sn-Pb2178composite system as a representative example with an experimental platform. Sn-Pb alloy is a class of solder alloysystem,11,12 and the ability to print these metal matrix composites will also offer insight into many applicationssuch as printed electronic interconnects and three dimensional electronic circuit architectures in space for thefuture.13Materials and MethodAlloy wires of 2mm diameter with a normal composition of Sn50Pb 50 (at. %) and carbon fibers were used asraw materials.

%) and carbon fibers were used asraw materials. One of the simplest methods of measuring the quantities for estimating the adhesion properties ofmaterials (i.e., the adhesion work, the surface energy, and the interfacial tension between certain liquids and asurface) requires the determination of the contact angle θ between the liquid and the surface.14 According to thepreliminary experimental results, the molten Sn 50Pb 50 cannot wet the carbon fiber. Hence, in order to improvethe adhesion property, carbon fibers were treated by electroplating two layers of metallic coating successively,1μm Ni coating which is used to improve the bonding strength between Cu coating and carbon fiber, and 2μm Cucoating. To identify the wetting behavior, a Sn 50Pb 50 ingot cube with 2mm side length was put on a plate coveredwith Cu/Ni coated carbon fibers in the vacuum furnace (Fig. 1(a)). The wetting angle was almost 109.6-degreewhen the temperature was 300℃ (Fig.

1(a)). The wetting angle was almost 109.6-degreewhen the temperature was 300℃ (Fig. 1(b)), and the phenomenon is non-wetting because of the oxidation of Cucoating. Oxygen contamination (even at ppm levels) can have a dramatic effect on the surface tension ofmetals.10 Hence the soldering flux was utilized to modify the Cu/Ni coated carbon fibers. The major ingredient ofthe soldering flux is rosin which eliminates the oxide layer and improves wetting behavior. The fibers werecovered with cream-state soldering flux homogeneously and it had a great influence (Figs. 1(c) and 1(d)).Based on the above wetting mechanism, FTP of Cf/Sn 50Pb 50 composites are carried out using a continuousfiber reinforced composites 3D printer (COMBOT-I from Shaanxi Fibertech Technology Development Co., Ltd,China). As the coated fibers are pulled out of the nozzle, solidification of metal matrix adhering to the fibersrapidly occurs, which forms homogeneous cylindrical composite filament (Figs.

As the coated fibers are pulled out of the nozzle, solidification of metal matrix adhering to the fibersrapidly occurs, which forms homogeneous cylindrical composite filament (Figs. 2(a) and 2(b)). In order to achievea precise feature size, the pulling rate should match to the solidification rate15. Therefore, the pulling rate in therange of 1-7mm/s is adopted according to the experimental results. When the pulling rate is large, the wetting isincomplete and the fiber is partial exposed. On the contrary, the uniformity decreases when the pulling rate is low.The temperature is in the range of 230-245℃.2179FIG. 1.

On the contrary, the uniformity decreases when the pulling rate is low.The temperature is in the range of 230-245℃.2179FIG. 1. Wetting angle test of Sn50Pb50 alloy on Cu/Ni coating carbon fibers (a) fibers without soldering flux at 25℃, (b) fibers withoutsoldering flux at 300℃, (c) fibers with soldering flux at 25℃, (d) fibers with soldering flux at 300℃Results and discussionThe approach of FTP is a dynamic wetting behavior of cylindrical surface in non-equilibrium condition.16,17Hence the wetting behavior of a fixed point on the fiber will change with the position and circumstance of themolten cavity.Two bunches of composite filaments were printed (Fig. 2 (b)) at 230℃ and 245℃ respectively with thespeed of 1mm/s. The tension strength of the 230℃ specimen is 60.4Mpa, the fracture surface (Fig. 2 (c)) showsthat the Cu/Ni coating still adheres to the fibers.

Toidentify the interfacial reactions of the printed C f/Sn 50Pb 50 composite, the microstructures of specimens printedat different pulling rates and the temperature of 235℃ were analyzed by the Gemini SEM 500 scanning electronmicroscope at Instrument Analysis Center of Xi'an Jiaotong University. The coating of Cu together with a part of Nidisappeared at the rate of 1mm/s, which formed an irregular edge (Fig. 3(a)). Improving the pulling rate to 3mm/scan keep the edge of Ni coating smooth and intact (Fig. 3(b)) and it is worth noting that the microstructuralmorphology of the two later prints show the existing of Cu coating at the rate of 5mm/s and 7mm/s (Fig. 3(c) and3(d) ).

3(b)) and it is worth noting that the microstructuralmorphology of the two later prints show the existing of Cu coating at the rate of 5mm/s and 7mm/s (Fig. 3(c) and3(d) ). That is mainly because a fast dissolution of Cu/Ni and interfacial reactions occurred during the slowprinting process.18,19 It is worth nothing that interfacial reactions between Cu and Sn formed the intermetalliccompounds, such as Cu 6Sn 518and Cu 3Sn19, which guarantee the reaction bonding in the interface. Hence FTPprovides a powerful tool to tailor the interfacial bonding properties of composites by controlling interfacialreactions. Furthermore, a frame which is 10 layers (0.5mm height per layer) was printed by this approach (Fig. 4).FIG. 4 The frame printed by FTP.2181ConclusionIn summary, an alternative approach is introduced in this paper to realize the additive manufacturing ofMMC.

4).FIG. 4 The frame printed by FTP.2181ConclusionIn summary, an alternative approach is introduced in this paper to realize the additive manufacturing ofMMC. The specimen at 245℃ has larger tension strength and the coatings of carbon fiber are separated becauseof the better adhesion property. The level of interfacial reactions decreases with increasing of printing rate from1mm/s to 7mm/s at 235℃. At the rate of 1mm/s and 3mm/s, the Cu coating is entirely consumed with theinterfacial reactions.The mechanisms of the FTP printed C f/Sn50Pb 50 prototype demonstrated in this study can be universallyextended to a vast range of other continuous fiber reinforced metal matrix composite system that possessappropriate wetting behavior. FTP of more technically interested metal matrix composite systems such as SiC f/Al,G f/Mg and SiC f/Ti will be a key subject of future work.Reference1. D.J. Korsmeyer, R.R. Landis, P.A.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFIBER-FED LASER-HEATED PROCESS FOR PRINTING TRANSPARENT GLASSJohn M. Hostetler i, Jonathan T. Goldstein ii, Douglas Bristow i, Robert Landers i, Edward C. Kinzel ii Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO iiAir Force Research Laboratory, Materials and Manufacturing Directorate, Wright-Patterson Air-Force Base, OHAbstractThis paper presents the Additive Manufacturing (AM) of glass using a fiber-fed process. Glass fiber with a diameter of 100 μm is fed into a laser generated melt pool. A CO2 laser beam is focused on the intersection between the fiber and the work piece which is positioned on a four- axis computer controlled stage. The laser energy at λ=10.6 μm is directly absorbed by the silica and locally heats the glass above the working point.

The laser energy at λ=10.6 μm is directly absorbed by the silica and locally heats the glass above the working point. By carefully controlling the laser power, scan speed, and feed rate, bubble free shapes can be deposited including trusses and basic lenses. Issues unique to the process are discussed, including the thermal breakdown of the glass, buckling of the fiber against an inadequately heated stiff molten region, and dimensional control when depositing viscous material.IntroductionAdditive manufacturing has revolutionized the field of manufacturing in recent years due to the advantages that these processes entail. Traditionally these advantages have pertained to an increased design freedom and the allowance of complicated three-dimensional geometries, rapid prototyping, more efficient material usage, and the capability to manufacture parts from functionally graded materials.

As these techniques for additive manufacturing mature however, more manufacturing applications become candidates for adaption these processes. Specifically, applications for optical systems are emerging as a promising enterprise; applications such as photonics packaging, gradient index (GRIN) and freeform optics, and integrated optics all stand to benefit from the development of a system which is capable of depositing high quality optical materials. This is due in part to the attractive material properties of glass itself. Glass is transparent in the visible spectrum, is amorphous and therefore does not suffer from grain boundary scattering, it is chemically inert, harder than transparent polymer counterparts, and displays a low sensitivity to temperature gradients.toWhile additive manufacturing has been widely studied for use with polymers, metals and ceramics, there is comparatively little work done in developing AM techniques for glass[1].

Those studies which have been conducted into the AM of glass share a limitation in depositing transparent glass due to bubble entrapment. Several groups utilizing selective laser sintering were able to fabricate three-dimensional glass parts[2,3,4], but these parts resemble sugar cubes in both texture and appearance. The development of continuous melting approaches for glass AM1594by several groups has been much more successful in depositing transparent glass. A filament-fed laser-heated process developed by Luo et al. demonstrated an ability to deposit void-free transparent glass using a feedstock with diameters ranging from 0.5-3 mm[5]. However, this process requires new cane to be reloaded into the filament feeder intermittently during the deposition process, and even the smallest diameter feedstock used of 0.5 mm cannot be rolled to extend the duration of uninterrupted deposition. Subsequently, Klein et al.

Subsequently, Klein et al. developed a gravity- fed molten vat approach[6] and Micron3dp introduced their high temperature extrusion technique[7]. Both of these platforms are capable of depositing very intricate three-dimensional geometries out of glass, but the inability of these techniques to deposit void-free blocks of glass constrains their use for optical applications.In light of the limitations with the filament-fed laser-heated process[5], a system utilizing optical fiber as a feedstock is very attractive, as kilometers of fiber may be spooled to allow for a continuous deposition process. In addition, the smaller diameter of optical fiber minimizes issues with thermal diffusion in the feedstock, potentially allowing for a greater volumetric deposition rate than that of the filament-fed process.

Optical fibers are widely available, consisting of extremely high quality, low loss glass (<1 dB/km), and are relatively cheap with prices lower than $180 per km of single mode fiber[8].Experiential ResultsThe experimental platform used in this study (Fig. 1) utilizes a CW CO2 laser (Synrad Evolution 125, λ0=10.6 μm, 140 µm spot size) which is incident on a soda-lime substrate. The substrate is fixed to a heater capable of reaching temperatures of 650℃, which is utilized to prevent thermal shock during the deposition process. The heater and substrate are in turn attached to a set of x-y-z stages, where the x and y stages (Aerotech ANT130-160XY) realize the horizontal movements, and a z stage (Aerotech ATS100-150) is used to move the platform upwards and downwards. A rotational stage (Aerotech ANT130-360-R) was installed on top of the x-y-z stages to enable rotations of the substrate.

A rotational stage (Aerotech ANT130-360-R) was installed on top of the x-y-z stages to enable rotations of the substrate. The fiber was fed into the melt pool (intersection of laser beam and substrate) by a custom-designed fiber feeder. 1% of the laser energy is reflected into a thermopile type power meter (Ophir 10A-V1) so that the laser power at the printing surface may be determined.Laser power, PCO2 laserBeam ShaperPyrometerSpectrometerDiode LaserPolarizerCameraBeam ExpanderFeed Rate, fShutterPower MeterScan Speed, vPrinted Part Substrate Heater Rotary Stage x-y-z stageFig. 1. Schematic of experimental platform1595Incandescent light is emitted from the melt pool during the printing process, with the spectrum of the radiation being dependent on the temperature of the melted glass.

This incandescent emission is collected using an OceanOptics USB-4000 fiber-coupled spectrometer (calibrated with an OceanOptics LS-1-CA 2800 K light source) which has a 0.8 mm diameter interrogation region centered on the laser heated area. The process parameters investigated in this study for their effect on the deposited glass are specified to be the laser power, P, fiber feed rate, f, and the platform scan speed, v (Fig. 2). The latter two parameters were consolidated into a dimensionless ratio of feed rate to scan speed, f/v.Pfv5 mmFig. 2. Process parameters for fiber-printing specified for their effect on the deposited glassThe fiber used in this study was 100 µm soda-lime fiber provided by Schott. One of the challenges in working with this fiber involves the necessity to remove the encapsulating sizing on the fiber intended to prevent the fiber from fracturing at microcracks.

One of the challenges in working with this fiber involves the necessity to remove the encapsulating sizing on the fiber intended to prevent the fiber from fracturing at microcracks. Fiber stripping tools are widely available commercially, but these tools are typically only used to remove sizing on small strips of fiber, and when it is attempted to strip the fiber continuously, these mechanical methods result in the accumulation of residual sizing particulate on the fiber itself which burn off and deposit soot, thus limiting the transmissivity of deposited glass.

Dissolving the sizing by soaking the fiber in ethanol for approximately thirty minutes prior to deposition proved to be a much more efficient approach, with initial transmissivity measurements finding that the transmissivity of the deposited fiber increased from 0.54 for untreated fiber to 0.79 after an ethanol bath.Another significant challenge in developing a fiber-fed process is the act of feeding compliant fiber into a viscoelastic molten region. This requires a fiber feeder capable of addressing both fiber deflection and fiber buckling. To prohibit the fiber from deflecting out of the melt pool, a series of nested hypodermic tubes were used to guide the fiber from the feed wheels to the deposition site. (Fig. 3). The inner diameter of this tubing was chosen to be 140 µm to limit the friction between the fiber and the inner walls of the tubing.CO2 laser beamFiber feedTelescoped hypodermic tubingMolten TrackFiberSubstrateFig. 3.

3. Schematic of telescoping hypodermic tubing used to prohibit fiber deflection1596The issue of fiber buckling continued to manifest in the unsupported span between the point of contact of the feed wheels and the entrance to the hypodermic assembly, labeled ‘d’ in Fig. 4. Here the fiber would buckle to such an extent as to cause it to snap, resulting in a deposition failure. It was found that this span, d, must be no greater than 2.4 mm in order to prevent the fiber from buckling. With these issues addressed, a continuous fiber deposition was enabled.dFig. 4. Fiber feeder with span, d, identified as critical parameter to prevent fiber bucklingThe process parameters P and f/v were first studied to identify their effects on the morphology of single tracks. Single tracks were deposited therefore over a range of parametric combinations of P and f/v, after which these tracks were cut and their cross sections polished.

Single tracks were deposited therefore over a range of parametric combinations of P and f/v, after which these tracks were cut and their cross sections polished. The track dimensions of height, h, width, w, and contact angle, θ, were then measured from these polished cross sections (Fig. 5).2.8 W3.1 W3.4 W3.9 W4.4 W4.8 Whθw1 mmFig. 5. Cross section of single tracks with measured dimensions labeled1806040h / w20f/v5:1 10:1 20:1] g e d [e g n Alt c a t n o C15012090f/v5:1 10:1 20:1600 2.53.03.5 Power [W]4.04.55.02.53.03.5 Power [W]4.04.5Fig. 6. Recorded dimensions of single track morphology as a function of process parameters15975.0The dimensions recorded of single track morphologies (Fig. 6) were then utilized to determine appropriate step sizes for depositing 2D walls, consisting of several stacked single tracks. These walls were deposited at representative parametric combinations of f/v and P.

These walls were deposited at representative parametric combinations of f/v and P. After deposition, these walls were polished on both sides to create transmissivity samples (Fig. 7).Fig. 7. 2D wall post-deposition (left), after polishing to create transmissivity sample (right)Newport power meter (1935-C)HeNe laser (λ = 633 nm)SampleFig. 8. Schematic of transmissivity measurement setupThe transmissivity of these polished samples were measured at a wavelength λ = 632.8 nm with a HeNe laser and a Newport power meter (1935-C). This was done by measuring the power of the HeNe laser beam unobstructed, then re-measuring the power of the beam after it is transmitted through a sample. It was found that the transmissivity trend of the deposited fiber was parabolic (Fig. 9). At lower laser powers, the fiber is softened and coils along the track, scattering any incident light and subsequently lowering the deposited part’s transmissivity.

9). At lower laser powers, the fiber is softened and coils along the track, scattering any incident light and subsequently lowering the deposited part’s transmissivity. At excessive laser powers, the soda-lime breaks down and is vaporized, depositing soot on the part which also lowers transmissivity. Between these two extreme cases however, a maximum transmissivity of 0.92 is measured for f/v = 15 and P = 6 W. This transmissivity value is ~93% of the transmissivity of conventionally manufactured soda-lime glass[9].]0.51 -m m[t n e c i f f e o C n o i t c n i t x Ei0.40.30.20.10.04.04.55.05.5 Power [W]6.06.5Fig. 9. Measured extinction coefficient of deposited fiber versus laser powerAfter determining the process parameters which result in the maximum transmissivity of the deposited soda-lime fiber, it was then attempted to deposit a simple lens with the fiber (Fig.159810).

The lens was printed in a single continuous spiral deposition from the center outwards, with a constant f/v to create a convex profile.Fig. 10. Simple lens printed with soda-lime fiber post-deposition: top view (left), side view (right)A large spot size (3 mm diameter) was used to reflow rings of the spiral during the deposition process to smoothen the lens surface; this reflow process however demonstrated a propensity for entrapping bubbles within the lens. Never the less, these simple printed lenses demonstrate the ability to optically focus light (Fig. 11).Fig. 11. Printed lens post-deposition demonstrating ability to focus lightIn addition to depositing lenses, the fiber-fed laser-melted process has been found to be capable of depositing complex free-standing 3D structures, such as the spiral shown in Fig. 12. These types of deposition are enabled by the temperature-dependent viscosity of soda-lime glass.

12. These types of deposition are enabled by the temperature-dependent viscosity of soda-lime glass. As the deposited fiber is moved away from the melt pool, it cools and the viscosity quickly increases, resulting in the structure becoming very rigid. There are several conditions however which must be maintained to ensure a successful deposition: In order to prevent the failure mode of ‘remelt’, wherein the laser softens a previously deposited structure causing it to collapse, the melt pool must be located at the focal point of the laser. In addition, an appropriate laser power must be used to ensure that the laser energy is completely absorbed by the melting fiber, and not allowed to continue beyond the deposition site and subsequently soften any preexisting structures beneath.1599Fig. 12. Free-standing spiral deposited with fiber-fed processConclusionsA fiber-fed laser-heated process is capable of depositing soda-lime glass parts with a 100 µm fiber feedstock.

12. Free-standing spiral deposited with fiber-fed processConclusionsA fiber-fed laser-heated process is capable of depositing soda-lime glass parts with a 100 µm fiber feedstock. Before deposition, an encapsulating sizing on the fiber must be removed via a thirty minute soak in ethanol. Telescoping hypodermic tubes successfully prevent the fiber from being deflected out of the melt pool, and the unsupported span between the entrance of these hypodermic tubes and the point of contact between the feed wheels must be less than 2.4 mm to prohibit fiber buckling. The fiber-fed process deposits glass with a maximum transmissivity of 0.92 (93% of the transmissivity of conventionally manufactured soda-lime glass) with the parameters f/v = 15 and P = 6 W.

The fiber-fed process deposits glass with a maximum transmissivity of 0.92 (93% of the transmissivity of conventionally manufactured soda-lime glass) with the parameters f/v = 15 and P = 6 W. The process has demonstrated an ability to deposit simple lenses capable of focusing light, as well as depositing complex free-standing three-dimensional structures.AcknowledgementsThis work was supported by the National Science Foundation (CMMI-1538464) as well as the Air Force Research Laboratory.References[1][2][3]G. Marchelli, R. Prabhakar, D. Storti, M. Gantor, “The guide to glass 3D printing: developments, methods, diagnostics and results”, Rapid Prototyping Journal 17(3), 187- 194 (2011). R. Khmyrov, S. Grigoriev, A. Okunkova, and A. Gusarov, "On the Possibility of Selective Laser Melting of Quartz Glass" Physics Procedia 56, 345-356 (2014). J.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFIBER-FED PRINTING OF FREE-FORM FREE-STANDING GLASS STRUCTURES John M. Hostetler i, Jason E. Johnson i, Jonathan T. Goldstein ii, Douglas Bristow i, Robert Landers i, Edward C. Kinzel ii Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO iiAir Force Research Laboratory, Materials and Manufacturing Directorate, Wright-Patterson Air-Force Base, OHAbstractAdditive Manufacturing (AM) of low-profile 2.5D glass structures has been demonstrated using a fiber-fed laser-heated process. In this process, glass fibers with diameters 90-125 µm are supported as they are fed into the intersection of the workpiece and a CO2 laser beam. The workpiece is positioned by a four-axis CNC stage with coordinated rotational/transitional kinematics.

The workpiece is positioned by a four-axis CNC stage with coordinated rotational/transitional kinematics. The laser energy at λ = 10.6 µm is coupled to phonon modes in the glass, locally heating it above its working point. The rapid heating and cooling process allows for the deposition of various glasses into free-standing three-dimensional structures such as trusses and other complex geometries. Issues unique to the process are discussed, including the thermal breakdown of the glass and index inhomogeneity between the fiber core and cladding when using single-mode optical fiber feedstock.IntroductionAdditive manufacturing has revolutionized the field of manufacturing in recent years due to the advantages that these processes entail.

Traditionally these advantages have pertained to an increased design freedom and the allowance of complicated three-dimensional geometries, rapid prototyping, more efficient material usage, and the capability to manufacture parts from functionally graded materials. As these techniques for additive manufacturing mature however, more manufacturing applications become candidates for adaption these processes. Specifically, applications for optical systems are emerging as a promising enterprise; applications such as photonics packaging, gradient index (GRIN) and freeform optics, and integrated optics all stand to benefit from the development of a system which is capable of depositing high quality optical materials. This is due in part to the attractive material properties of glass itself.

This is due in part to the attractive material properties of glass itself. Glass is transparent in the visible spectrum, is amorphous and therefore does not suffer from grain boundary scattering, it is chemically inert, harder than transparent polymer counterparts, and displays a low sensitivity to temperature gradients.toWhile additive manufacturing has been widely studied for use with polymers, metals and ceramics, there is comparatively little work done in developing AM techniques for glass[1]. Those studies which have been conducted into the AM of glass share a limitation in depositing transparent glass due to bubble entrapment. Several groups utilizing selective laser sintering were able to fabricate three-dimensional glass parts[2,3,4], but these parts resemble sugar cubes in both texture and appearance. The development of continuous melting approaches for glass AM994by several groups has been much more successful in depositing transparent glass.

The development of continuous melting approaches for glass AM994by several groups has been much more successful in depositing transparent glass. A filament-fed laser-heated process developed by Luo et al. demonstrated an ability to deposit void-free transparent glass using a feedstock with diameters ranging from 0.5-3 mm[7]. However, this process requires new cane to be reloaded into the filament feeder intermittently during the deposition process, and even the smallest diameter feedstock used of 0.5 mm cannot be rolled to extend the duration of uninterrupted deposition. Klein et al. developed a gravity-fed molten vat approach[5], after which Micron3dp introduced their high temperature extrusion technique[6].

Klein et al. developed a gravity-fed molten vat approach[5], after which Micron3dp introduced their high temperature extrusion technique[6]. Both of these platforms are capable of depositing very intricate three-dimensional geometries out of glass, but the inability of these techniques to deposit void-free blocks of glass constrains their use for optical applications.In light of the limitations with the filament-fed laser-heated process[7], a system utilizing optical fiber as a feedstock is very attractive, as kilometers of fiber may be spooled to allow for a continuous deposition process. In addition, the smaller diameter of optical fiber minimizes issues with thermal diffusion in the feedstock, potentially allowing for a greater volumetric deposition rate than that of the filament-fed process.

Optical fibers are widely available, consisting of extremely high quality, low loss glass (<1 dB/km), and are relatively cheap with a prices lower than $180 per km of single mode fiber[8].Experiential ResultsThe experimental platform used in this study (Fig. 1) utilizes a CW CO2 laser (Synrad Evolution 125, λ0=10.6 μm, 140 µm spot size) which is incident on a soda-lime substrate. The substrate is fixed to a heater capable of reaching temperatures of 650℃, which is utilized to prevent thermal shock during the deposition process. The heater and substrate are in turn attached to a set of x-y-z stages, where the x and y stages (Aerotech ANT130-160XY) realize the horizontal movements, and a z stage (Aerotech ATS100-150) is used to move the platform upwards and downwards. A rotational stage (Aerotech ANT130-360-R) was installed on top of the x-y-z stages to enable rotations of the substrate.

A rotational stage (Aerotech ANT130-360-R) was installed on top of the x-y-z stages to enable rotations of the substrate. The fiber was fed into the melt pool (intersection of laser beam and substrate) by a custom-designed fiber feeder. 1% of the laser energy is reflected into a thermopile type power meter (Ophir 10A-V1) so that the laser power at the printing surface may be determined. Incandescent light is emitted from the melt pool during the printing process, with the spectrum of the radiation being dependent on the temperature of the melted glass. This incandescent emission is collected using an OceanOptics USB-4000 fiber- coupled spectrometer (calibrated with an OceanOptics LS-1-CA 2800 K light source) which has a 0.8 mm diameter interrogation region centered on the laser heated area.995Fig. 1.

1. Schematic of experimental platformThe process parameters investigated in this study for their effect on the deposited glass are specified to be the laser power, P, fiber feed rate, f, platform scan speed, v, and the laser spot size (Fig. 2). The latter two parameters were consolidated into a dimensionless ratio of feed rate to scan speed, f/v. The feedstock used was SMF-28 optical fiber manufactured by Corning. This fiber has an acrylic coating which is mechanically stripped following a two-hour soak in denatured alcohol, leaving a cladded quartz core which has an outer diameter of 125 µm. The cladding has a refractive index of 1.44681, while the 9 µm diameter core has a refractive index of 1.45205.Fig. 2.

The cladding has a refractive index of 1.44681, while the 9 µm diameter core has a refractive index of 1.45205.Fig. 2. Process parameters for fiber-printing specified for their effect on the deposited glassWith the ability of the fiber-fed laser-heated process to deposit transparent quartz structures proven in previous studies, it was now attempted to deposit a spherical lens in such a way that the shape of the deposited lens adhered to a design profile (Fig. 3). Each layer consists of a series of concentric circles, starting from the outside and working its way inwards. Each ring is designed to maintain a constant feed rate-to-scan speed ratio of 1:1. After depositing each layer, the CO2 laser is then defocused, and the layer is reflowed to create a smooth surface.996Fig. 3.

After depositing each layer, the CO2 laser is then defocused, and the layer is reflowed to create a smooth surface.996Fig. 3. Cross-Section of Lens Profile Design (top) Progression of Lens Deposition (bottom)These lenses deposited with Quartz SMF do an adequate job of imaging, be it the pixels on an LCD monitor, or an AFRL Resolution test chart (Figure 4). It is unclear at this time whether or not the quartz glass within the core is mixing with the cladding during the deposition, or if there remains segregation between the slightly higher index of refraction glass in the core, and that of the cladding. Any effects of this index inhomogeneity are soon to be measured and compared with lenses which are deposited using specially drawn quartz fiber, where the entire fiber is comprised of material with the same index of refraction.Fig. 4.

4. Profile of spherical lens (top left), lens imaging pixels on LCD monitor (bottom left), AFRL resolution test chart with no lens (top right), AFRL resolution test chart imaged through spherical lens (bottom right)The fiber-fed process is not limited however to only printing transmissive parts; by printing with a feed rate-to-scan speed ratio of 1:1 and a small spot size of ~200 µm, large 3D parts may be printed with quartz SMF fiber. Simple geometries such as those depicted in figure 5 have997been deposited with heights of up to 15 mm, and more complex geometries such as arches (Fig. 6) have been demonstrated with parts successfully deposited up to 20 mm high. The more complex geometries such as that depicted in figure 11 have been made possible by running a 3D model through a slicer program, thus discretizing the shape into a series of points. Then, a G- code generator is used to produce motion commands to move form one point to the next.

Then, a G- code generator is used to produce motion commands to move form one point to the next. However, the path planning often requires manual input to select the next appropriate vector when moving from one z-position to the next.Fig. 5. Simple geometries deposited with quartz SMFFig. 6. St. Louis Gateway Arch generated deposition path (left) structure post-deposition (right)In addition to the continuous structures discussed above, more free-standing complicated geometries involving many starts and stops in the deposition have been accomplished, such as trusses (Fig. 7). These trusses involve the deposition of free-standing fiber in directions which are not aligned with the fiber feeder direction, a disconnect between the substrate and the fiber998feeder, and a subsequent reconnection for the next sequence in the deposition.

The necessity for consistent disconnects and reconnects which do not damage the already deposited structure requires the use of lead-ins and lead-outs, which often result in unnecessary segments of fiber deposited along the periphery of the truss. The possibility of adding a terminal portion to the deposition code which uses the laser to sever the connections between the intended truss structure and unnecessary segments of fiber deposited to create a lead-in or lead-out will soon be tested.Fig. 7. Fiber truss in various stages of depositionFinally, it has been demonstrated that with the laser-heated fiber deposition process, quartz SMF fiber may be deposited which is capable of guiding a HeNE laser beam (Fig. 8). A segment of the fiber is deposited in free space, and then is sloped downwards to connect to the substrate while rotating 90 degrees, and the remainder is deposited along the substrate surface.

8). A segment of the fiber is deposited in free space, and then is sloped downwards to connect to the substrate while rotating 90 degrees, and the remainder is deposited along the substrate surface. This is done to ensure that any light that appears to be guided by the deposited fiber is done so because the HeNe beam is in fact coupled within the fiber itself, and not simply reflection within the substrate.Fig. 8. HeNe laser beam guided through deposited quartz SMF fiberAs proof that these fibers do in fact guide light from one end of the fiber to the other, and that any light coupled into the substrate does not interfere with any signaling, several waveguides were deposited next to each other and illuminated with the HeNe laser individually (Fig. 9).999Fig. 9.

9).999Fig. 9. Four adjacent waveguides demonstrating ability to isolate waveguide illuminationFigure 9 not only shows that the light can be transmitted around a circular arc with a radius of 1mm, but also that each fiber may be illuminated individually. It may be seen in figure 9 that there are some instances where there is a slight illumination in adjacent waveguides, as well as from the substrate as well. This is believed to be due to two factors, the first being dust and other particulate that accumulate on the substrate surface, which scatters light. The second, much less trivial, is the coupling of the laser beam from the waveguide into the substrate. This happens at the spot where the fiber begins to lift off the substrate surface to slope upwards to the diagonal section. This lift off is done at a sharp angle, which is undesirable for waveguides; This is a coding limitation for this platform at this time.

This lift off is done at a sharp angle, which is undesirable for waveguides; This is a coding limitation for this platform at this time. There is currently no way to arc the stage in the Z direction, and so the sharp angle scatters light and couples a large percentage of the beam into the substrate. This is a temporary limitation however, as a new program will be introduced soon which will allow for arcing the Z-stage along the X or Y direction.It was also observed that when the quartz SMF fiber is deposited in free space, certain portions of the fiber will scatter light. This is clearly an undesirable characteristic for a waveguide, so the causes behind this phenomenon are being investigated. Sections of the free- standing fiber which scatter light were observed under a microscope, revealing that these areas appear to be necking, while segments which did not scatter light have a more uniform cross- section (Fig. 10).

10). At this time it is speculated that the fiber necking is due to small interruptions in the fiber feed, due either to slipping in the feed wheels or backlash in the bevel gears used to drive the feed wheels. A new fiber feeder is currently being designed therefore which will address these issues and hopefully prevent them in the future.1000Fig. 10. Freestanding fiber observed under a microscope: Necking site which scatters light (top right), uniform segment which does not scatter light (bottom right)ConclusionsThe fiber-fed laser-heated process is capable of depositing quartz SMF that has a nominal diameter of 125 µm, where f=v=1 mm/s. The fiber-fed laser-heated process was used to deposit simple spherical lenses which demonstrated the ability to optically focus light. The deposition approach for the spherical lens was improved by depositing the lens one layer at a time, where each layer consists of concentric rings deposited with a feed rate-to-scan speed ratio of 1:1.

The fiber-fed process was then used to print three-dimensional structures ranging from complex continuous shapes such as the St. Louis Gateway Arch, to discontinuous free-standing structures such as simple trusses. Lastly, it was demonstrated that the deposited quartz fiber is able guide light, prompting further investigations into the utility of using this process for printing waveguides in the future.AcknowledgementsThis work was supported by the National Science Foundation (CMMI-1538464) as well as the Air Force Research Laboratory.References[1][2][3]G. Marchelli, R. Prabhakar, D. Storti, M. Gantor, “The guide to glass 3D printing: developments, methods, diagnostics and results”, Rapid Prototyping Journal 17/3 187- 194 (2011). R.S. Khmyrov, S.N. Grigoriev, A.A. Okunkova, and A.V. Gusarov, "On the Possibility of Selective Laser Melting of Quartz Glass" Physics Procedia 56 345-356 (2014). J.

As the demand for sand has tripled over the last two decades, finding new solutions for thermostable binder systems and ensuring the efficient use of resources is essential. To address this issue, the REINFORCED SAND project is exploring glass fiber-reinforced sand molds and temperature-stable inorganic binders within the binder-jetting process. The aim is to improve the mechanical properties of the (cid:28)D-printed casting molds to make multiple uses of these possible, reducing resource consumption and increasing profitability. For this purpose, various material systems and manufacturing parameters were examined. It has been demonstrated that fiber-reinforced sand molds can be used for multiple castings, and fibers can be principally processed in the binder jetting process. These efforts aim to make sand casting production more environmentally friendly and sustainable.

These efforts aim to make sand casting production more environmentally friendly and sustainable. The REINFORCED SAND project is a step towards achieving these goals.IntroductionThe casting technology, compared to other manufacturing processes, enables cost-efficient production of highly geometrically complex and function-integrated parts obeying the following process steps:The mold halves and cores define the cast. • The casting mold is assembled with these mold halves and cores. • Subsequently, the part is cast, cooled down, and removed from the mold. • Finally, the cores are removed with the help of mechanical or chemical processes and machined if necessary.Mass production with a large capacity typically involves the use of permanent molds, usually made from metal. Prototypes and small to middle production mostly use the sand casting process, where the molds and cores are only used once [(cid:6)].

Prototypes and small to middle production mostly use the sand casting process, where the molds and cores are only used once [(cid:6)]. This leads to high resource consumption and a challenging recycling process due to organic or inorganic sand mold additives [ ]. The demand for sand has tripled to (cid:4)! to (cid:2)! billion metric tons per year in the past two decades, and sand is the second most traded resource in the world after water [(cid:28)]. Despite the large resource deposit, it is desirable to guarantee resource-saving via further processing and reprocessing. To create an environmentally friendly and sustained production chain in sand casting, new solutions are required for the binder systems and to integrate an efficient use of resources through reusable casting molds.The aim of reinforcing sand molds for the casting industry with fibers is to find a middle ground between the immense investment for permanent molds and the single-use of sand molds.

Incorporating fibers into the sand molds is anticipated to improve the mechanical characteristics of the disposable molds and cores, preventing the degradation of the fiber-reinforced molds, and enabling their reuse.Furthermore, sand cores enable the production of highly complex internal structures. These cores also consist of a mold base material and a binder that connects the mold base material. Due to their micromechanical structure, sand cores exhibit highly brittle failure behavior. Because of this behavior, sand cores hardly undergo any plastic deformation and disintegrate immediately.

Due to their micromechanical structure, sand cores exhibit highly brittle failure behavior. Because of this behavior, sand cores hardly undergo any plastic deformation and disintegrate immediately. This can particularly be a challenge for handling processes, as the sand cores fail brittly when overloaded and can bring automated production to a hault, as the brittly failed sand cores disintegrate into many fragments and cannot be removed as a whole from the process chain by a robot when a defect is detected in the core.1600The fibers in the molding material that absorb the deformation energy may shift the highly brittle failure behavior of sand molds and cores into a ductile material behavior. Further, an inorganic binder is used in order to minimize harmful combustion gases to the environment and health that emerge from decomposing organic binder i.e.

Further, an inorganic binder is used in order to minimize harmful combustion gases to the environment and health that emerge from decomposing organic binder i.e. furan and phenolic resin [(cid:4)].State of the ArtSand Cores and Sand Forms in Casting IndustrySand cores are often used to produce highly complex internal structures in cast components. These are usually made from a mold base material such as silica sand and a binder with chemical additives. Due to increased environmental requirements, inorganic binders based on water glass are increasingly being used. In the industrial environment, sand cores are produced either by core shooting or binder jetting.

In the industrial environment, sand cores are produced either by core shooting or binder jetting. In both processes, the mold base material is mixed with binder and additives and, in the case of inorganic binder systems, the sand-binder system is then solidified by drying the liquid binder phase.Mechanical Properties of Sand CoresIn industry, minimum tensile strengths between .(cid:3) and (cid:28).(cid:28) MPa are usually required for sand cores [(cid:2)]. Due to their structure, sand cores have a porous microstructure and can be classified as granular media. These exhibit a pressure-dependent material behavior and thus endure significantly more compressive loads compared to tensile loads. The mechanical properties of sand cores are often characterized in the foundry industry using bending bars [(cid:2),(cid:5)]. The bending strength was measured from breaking the bars in bending tests. Sand cores exhibit brittle failure and typical elastic moduli between (cid:5) and (cid:1) GPa.

The bending strength was measured from breaking the bars in bending tests. Sand cores exhibit brittle failure and typical elastic moduli between (cid:5) and (cid:1) GPa. Lechner et. al. [(cid:3)] were able to demonstrate the highly brittle strain behavior with microstructural cracking before macroscopic visibility optically.Fiber Reinforcement of powder based materials in different fields of useFiber reinforcement of brittle materials is also used in the field of construction materials. In this process, the properties of concrete are improved by adding fibers made of glass, minerals, metals or polymers [(cid:1)]. Fiber- reinforced concrete (FRC) has improved crack resistance and energy absorption capacity compared to normal concrete [(cid:7)]. This is practically used in the field of earthquake resistance [(cid:6)!] or splinter protection in the event of explosions [(cid:6)(cid:6)] or fire [(cid:6) ].

This is practically used in the field of earthquake resistance [(cid:6)!] or splinter protection in the event of explosions [(cid:6)(cid:6)] or fire [(cid:6) ]. Shrinkage cracks occurring during the curing of concrete can also be reduced by fiber reinforcement [(cid:6)(cid:28)].In all the aforementioned applications, fiber reinforcement achieves only a slight improvement in strength before the first cracks occur, i.e., in the first part of the stress-strain curve up to the first maximum, there is only a slight difference compared to unreinforced concrete. The benefit of fiber-reinforced concrete becomes evident after the initial crack forms. Unlike purely brittle materials that abruptly lose strength, the fibers help hold the material together even after cracking, allowing the force to continue being transferred. This property is similar to the behavior of ductile materials, in which plastic deformation occurs when overstressed, but the material is not spontaneously destroyed.

This property is similar to the behavior of ductile materials, in which plastic deformation occurs when overstressed, but the material is not spontaneously destroyed. For this reason, we will refer to this fracture behavior as quasi-ductile throughout the paper. Eldesouky et. al. [(cid:6)(cid:4)] showed that randomly distributed fiber reinforcement of dry sand can increase the shear strength and dilation of sand, which could be used for soil stabilization. An investigation was conducted on the use of fiber reinforcement with polypropylene fiber, aluminum silicate fiber, and glass fiber in a liquid water glass binder for investment casting. To increase the strength of inorganically bonded investment casting molds, Lü et al. [(cid:6)(cid:2)] mixed various fibers into the mold material and determined the green strength, firing strength, and residual strength of these fiber-reinforced shells.

KG) with the addition of liquid water glass binder (IOB binder, voxeljet AG) at a speed between (cid:28)(cid:2)! and (cid:2)!! rpm for a maximum of five minutes. This speed was chosen based on experience, as the best mixing was observed at these speeds.Different glass fiber (GF) types (overview in Table (cid:8)) and lengths are mixed together with the mold base material and the liquid water glass binder until all components are wetted as well as possible. The goal was to achieve a homogeneous mixture of fiber-binder mixture with the added fibers.GF 1 (alkali resistant glass fiber designed for concrete applications) and GF 6 (glass fiber designed for gypsum or wet-laid processes) are glass fibers commonly used in inorganic binder systems, which is why they were chosen.

GF 2 and GF 3, both glass fibers designed for use with organic binder resins, were also tested using the liquid water glass binder.To test the influence of fiber surface coating on the miscibility of fibers in the sand-binder system, an uncoated glass fiber GF 4 was chosen. The uncoated glass fiber used is a standard alkali-free E-glass fiber with a diameter of 16 µm, which is one of the most widely used types of glass fiber worldwide.To reach the lower limit of the fiber length range, ground fibers with a length of !! µm were also tested. To test the recyclability the recycled E-glass fiber GF (cid:3) with a diameter of about (cid:6)(cid:2) µm was chosen. It is coming from a two-step pyrolysis process. The first step is a pyrolysis step under nitrogen atmosphere and the second step an oxidation step under synthetic air with ! % oxygen to remove char from the surface of the fiber. The fibers were reclaimed from an UP/PMMA-resin composite at (cid:2)!! – (cid:2)(cid:2)! °C.

The fibers were reclaimed from an UP/PMMA-resin composite at (cid:2)!! – (cid:2)(cid:2)! °C. This fiber was chosen to investigate the influence of the recycling process on the mechanical properties of the fiber-reinforced material.The mixing time required for a homogeneous mixture is indicated in Table .Figure (cid:8): Process chain for specimen productionFor every bending bar approximately (cid:6)(cid:28)(cid:2) g of sand-binder-fiber mixture was stamped into a steel mold and afterwards cured in a furnace at (cid:1)! °C for (cid:6).(cid:2) h, except for GF (cid:3), where !! °C was used according to [(cid:6)(cid:1)].

Five bending bars were produced for each experiment.Table ,: Experimental design.Experiment number GF1_1 GF2_1 GF2_2 GF2_3 GF2_4 GF2_5 GF2_6 GF2_7 GF2_8 GF2_9 GF2_10 GF3_1 GF4_1 GF4_2 GF4_3 GF5_1 GF5_2 GF5_3 GF6_1 GF6_2 GF7_1 GF7_2Fiber typeGF 1 GF 2 GF 2 GF 2 GF 2 GF 2 GF 2 GF 2 GF 2 GF 2 GF 2 GF 3 GF 4 GF 4 GF 4 GF 5 GF 5 GF 5 GF 6 GF 6 GF 7 GF 7Fiber length in mm 6.3 6.3 6.3 6.3 6.3 6.3 6.3 4.2 16.8 6.3 6.3 6.3 16.8 4.2 6.3 0.2 0.2 0.2 6 6 6 12Amount of fibers in wt.% 0.5 0.5 1 2.5 3.75 5 7 2.5 2.5 2.5 2.5 5 2.5 2.5 5 1 2 5 2.5 5 5 5Amount of Binder in wt.% 3.5 3.5 3.5 3.5 3.5 3.5 3.5 3.5 3.5 7 14 3.5 3.5 3.5 3.5 3.5 3.5 3.5 3.5 3.5 3.5 3.5Mixing Time in minutes 15 15 15 15 15 15 15 15 101 15 15 15 101 15 15 5 5 5 5 5 5 51 Manually mixed into the sand-binder mixture.1603Mold Production and CastingTo test the fiber reinforced molding material in the casting process, a mold in the shape of a knight chess figure – as can be seen in Figure 2 – was selected.

This shape was chosen because of its sophisticated shape and the amount of detail that allows to make a statement about the surface quality. The maximum height of the figure from the socket to the top of the head is 75 mm, and it has a maximum width of 50 mm. The mold base material was stirred in a mixer with a whisk attachment with the addition of liquid water glass binder at a speed level three for (cid:5)! seconds. GF 2, the glass fiber type that yielded the best results in the “Specimen Production” chapter, was continuously added in small amounts within 1 minute of continuous stirring to prevent fiber agglomeration. Subsequently, the sand‑binder‑fiber system was pressed into an aluminum mold with the positive chess‑figure model fixed in the mold. Afterwards, the mold was cured in an oven at !! °C for two hours.

Subsequently, the sand‑binder‑fiber system was pressed into an aluminum mold with the positive chess‑figure model fixed in the mold. Afterwards, the mold was cured in an oven at !! °C for two hours. In parallel, a reference casting mold without any fiber content was also produced for each fiber‑reinforced casting mold with different fiber ratios to ensure comparability. In addition to each mold, five bending bars – as described in the chapter "Specimen Production" – were made from the same batch for comparison purposes. Three different GF fiber fractions, each with (cid:28) wt.%, (cid:2) wt.%, and (cid:3) wt.% were investigated.The molds were then cast with a standard AlSi-alloy at (cid:5)(cid:1)! °C and tested for reusability. If the mold had not been destroyed during casting and demolding, another casting attempt was made.

If the mold had not been destroyed during casting and demolding, another casting attempt was made. This process was repeated until the mold was destroyed.Figure 2: Aluminum mold with a model of a knight chess figure for the production of sand casting molds.Binder-Jetting of Fiber-reinforced Bending BarsTo test the general processability of glass fibers in (cid:28)D-printing, bending bars – as described in the chapter “Specimen production” – were produced on two different (cid:28)D-printing systems. In these experiments, the processability of the GF fibers was the primary focus; hence, the quantity of fibers introduced had not been examined initially. Bending bars were printed along the coating direction. The coating mechanism and binder system were investigated in relation to fiber reinforcement. The production of bending bars using the Binder Jetting method involves the following process steps: first, a layer of sand is deposited and leveled.

The production of bending bars using the Binder Jetting method involves the following process steps: first, a layer of sand is deposited and leveled. Then, a binder agent is printed onto the leveled sand layer. The material solidifies at the locations where binder material is applied in a subsequent chemical step. The fibers were evenly spread manually using a (cid:6) mm mesh sieve for (cid:2) seconds per printed bending bar layer at a height of approximately (cid:6)(cid:2) cm. The bending bars were placed (cid:2) cm apart along the long side in the transverse direction within the build chamber. The process starts again, and sand is applied onto the printed and fiber‑scattered layer. This process is repeated layer by layer, creating a sand bed in which the desired object is located. After the curing process, the component is removed from the sand bed with full strength.

This process is repeated layer by layer, creating a sand bed in which the desired object is located. After the curing process, the component is removed from the sand bed with full strength. This process is schematically shown in Figure -.Coating a sand-fiber mixture proved to be impossible in a preliminary test. Firstly, due to electrostatic charging, fiber can hardly be mixed with dry sand. Secondly, the fibers block the narrow coating gap of the coater, making homogenous coating impossible.1604Figure -: Process of the Binder-Jetting method. The process steps “Coating” and “Printing” are repeated until the desired object is completely printed.Two binder systems were tested. The first one is a two-component system consisting of furan resin and an activator (RPT , voxeljet AG) – an acid – which is mixed into the sand prior to (cid:28)D printing. In this system, (cid:28) wt.% of furan resin was printed into the sand with !.(cid:28) wt.% content of activator.

In this system, (cid:28) wt.% of furan resin was printed into the sand with !.(cid:28) wt.% content of activator. Using the binder system, four bending bars (two with fiber reinforcement and two reference bars without fiber addition) were printed on a prototype in- house research (cid:28)D printer with a layer height of (cid:28)(cid:2)! micrometers in the coating direction. In this case, the coating module consists of a sand dispenser that sprinkles sand onto the previous layer and a counter rotating roller, smoothing and compressing the previously deposited sand layer. This process is shown in the schematic diagram of Figure 3.The second binder, a single component water glass (IOB, voxeljet AG), was used in the voxeljet VX(cid:6)!!! printer with a (cid:28) wt.% binder content.

Using this binder system, six bending bars (three with fiber reinforcement and three reference bending bars without fiber addition) were printed with a layer height of (cid:28)(cid:2)! micrometers in the coating direction. The coating module consists of a blade that smoothes and compresses the previously applied sand and a sand dispenser located directly above the blade. In this case, sand is not sprinkled onto the glass fibers but pushed over the fiber layer.Figure 3: Schematic diagram of the -D-printing process with help of a roller coater, used for the organic binder system.(cid:20)-Point-Bending TestAll specimens are tested by a (cid:28)-point-bending test on a universal testing machine (Z!(cid:2)!, ZwickRoell GmbH & Co. KG) with a (cid:6) kN load cell. The support has a distance of (cid:6)(cid:2)! mm, and the load is induced in the middle of the bending bar. Both the supports and the stamp have radii of (cid:2) mm.

KG) with a (cid:6) kN load cell. The support has a distance of (cid:6)(cid:2)! mm, and the load is induced in the middle of the bending bar. Both the supports and the stamp have radii of (cid:2) mm. The test is carried out with a test speed of (cid:2) mm/min. On the bottom of the specimen, a length gauge is installed to measure the bending distance of the specimen ((cid:28)-P bending feeler (cid:6)(cid:6) mm with makroXtens).

With the (cid:28)-point-bending test, the strength of the specimens as well as the elongation at break is measured, shown in Figure (cid:2), where F is the applied force, L-P is the distance between the two supports, I-P is the length of the specimen, h-P is the height, and w-P the width of the specimen.1605Figure :: Schematic diagram of the --point-bending test [(cid:8)<].GOMThe GOM ATOS Compact Scan (Carl Zeiss GOM Metrology GmbH) with the MV(cid:6)(cid:3)! attachment wasused to determine and compare the castings' surface with a measuring volume of (cid:6)(cid:3)! x (cid:6)(cid:28)! mm², a camera angle of (cid:3).(cid:6)°, and a camera resolution of (cid:6) megapixels.The open-source software CloudCompare (cloudcompare.org) was used to generate a point cloud and calculate the average absolute deviation (AAD) between the castings and the ideal CAD model.

The AAD was calculated using Equation (cid:6), where n is the number of measured points, and xi is the deviation from the measured point to its corresponding point with the minimum distance on the ideal CAD model. The first, fifth, and tenth castings were compared, produced using the mold with (cid:2) wt.% and (cid:3) wt.% GF , along with their respective reference castings.(cid:1) (cid:2)(cid:2)(cid:3)|(cid:5)(cid:6)| (cid:6)(cid:7)(cid:1)Equation (cid:8)ResultsSpecimen ProductionThe specimens were produced with different glass fibers, different fiber lengths and different amount of glass fiber and binder in the mixture and were tested with a (cid:28)-point-bending test. Here, the bending strength and the deflection were evaluated and are shown in the following Table -.

Here, the bending strength and the deflection were evaluated and are shown in the following Table -. It was found that some glass fibers and their blends with sand and binder result in much higher elongation at break at maximum force than the elongation at break at maximum force without glass fiber reinforcement. The breaking behavior changed from brittle behavior to a more quasi-ductile breaking behavior, see Figure @.] a P M[ σ5 4,5 4 3,5 3 2,5 2 1,5 1 0,5 0Sample with 5 wt% GF 2Reference sample0,4 ɛ [%] Figure @: Example of breaking behavior of a sample without glass fiber reinforcement and a sample of GF , with : wt.% of glass fiber. It can be seen that the unreinforced specimen (dotted line) shows a brittle fracture behavior, while the stress-strain curve of the fiber-reinforced specimen (solid line) resembles that of a ductile material.00,20,60,81606Table -: Results and standard deviation of hand made specimen for fiber type evaluation.

Only the mixture containing (cid:6)(cid:4) wt.% binder and .(cid:2) wt.% glass fibers resulted in a higher elongation at Fmax, which can be seen in Figure <.109] a P M87[ h t g n e r t s g n d n e Bi654321Bending strength of referenceBending strength with 2,5% GF 2Linear (Bending strength of reference ) Linear (Bending strength with 2,5% GF 2)0051015Amount of Binder [%]Figure B: Influence of binder content on bending strength with and without GF ,.16080,160,14]%[ k a e r b t a n o i t a g n o el0,120,100,080,060,040,020,00elongation at break of referenceelongation at break with 2,5% GF 2Linear (elongation at break of reference) Linear (elongation at break with 2,5% GF 2)051015Amount of Binder [%]Figure <: Influence of binder amount on elongation at break with and without GF ,.Casting of MoldsDifferent fiber ratios within the casting mold obtained varying numbers of castings with different surface qualities.

The criterion of failure was determined by considering the mold halves that were broken and could not be reused due to the casting process. At a fiber ratio of (cid:28) wt.%, the casting molds were completely destroyed after the fourth casting. At fiber ratios of (cid:2) wt.% and (cid:3) wt.%, ten castings were obtained for each mold. However, the casting attempts were discontinued due to poor surface quality and not due to complete destruction of the casting mold.

However, the casting attempts were discontinued due to poor surface quality and not due to complete destruction of the casting mold. The reference molds without fiber reinforcement – produced in parallel to their respective fiber-reinforced casting molds – always broke after the first casting.Table 3: Correlation between the fiber content within the casting mold and the achieved number of casts.Fiber content within the casting mold & wt.% (cid:20) wt.% ! wt.% # wt.%Achieved number of casts 1 4 10 10The bending bars, produced in parallel to the casting molds, were tested in a (cid:28)‑point bending test, as described in the “(cid:28)‑Point‑Bending Test” chapter, to establish a correlation with the material's mechanical properties.

Figure (cid:8)C shows the evaluated results of the (cid:28)-point bending tests conducted on the bending bars produced in parallel to the casting molds containing different fiber amounts.1609Figure (cid:8)C: --point-bending test results of the respective five bending bars produced in parallel to the casting molds. The fiber-reinforced (varying fiber content - %, : % and E %) and reference (C % fiber content) bending bars are compared.During the measurement of the surface deviation of the castings from the original CAD file, only the casting series with 5 wt.% and 7 wt.% glass fiber content were considered because, in the casting experiment, a minimum of 10 castings could be produced in these series (Table 3). The results of this measurement are shown in Figure (cid:8)(cid:8).1,401,201,161,000,86]m m[D A A0,800,600,400,410,550,250,350,520,645 wt.% GF7 wt.% GF0,200,00ref.

cast1st cast5th cast10th castFigure (cid:8)(cid:8): AAD6 in relation to the number of castings. The reference castings (ref. cast) do not contain any fiber content and were depicted for the sake of comparison.Multiple casting attempts were made per fiber content, however only one mold was used per fiber content. Therefore, only the trend can be observed, but the result is not statistically validated. The images generated using CloudCompare, which display the deviations of the individual casting from the designed form visually, are shown in Figure (cid:8),.6 Absolute Average Deviation1610CastCasts made from fiber- reinforced mold with a fiber content of ! wt.% (GF (cid:31))Casts made from fiber- reinforced mold with a fiber content of # wt.% (GF (cid:31))Ref. cast(cid:6)st cast(cid:2)th cast(cid:6)!th castFigure (cid:8),: The deviations of the castings from the CAD file can be easily observed.

cast(cid:6)st cast(cid:2)th cast(cid:6)!th castFigure (cid:8),: The deviations of the castings from the CAD file can be easily observed. With each casting attempt, the mold becomes more hollowed out, resulting in a greater deviation of the casting from the CAD file.1611Binder-Jetting of Fiber-reinforced Bending BarsThe bending bars, which were produced using different binder systems and coating methods in the (cid:28)D-printing process, were tested in the (cid:28)-point bending test. The mechanical properties investigated included bending stress and elongation at break at maximum bending stress of the material system. In the organic (cid:28)D- printing experiment, two bending bars were produced, while in the inorganic (cid:28)D-printing experiment, three bending bars were produced. The evaluated test results are shown in Figure (cid:8)-.Figure (cid:8)-: The left diagram shows the elongation at break at maximum stress.

The evaluated test results are shown in Figure (cid:8)-.Figure (cid:8)-: The left diagram shows the elongation at break at maximum stress. In the right diagram, the maximum stress is shown.An image of the side view of a (cid:28)D-printed bending bar, manufactured using the organic binder system (RPT , voxeljet AG), GS(cid:6)(cid:4)RP sand (Strobel Quarzsand GmbH), and (cid:5).(cid:28) mm long GF glass fibers, can be seen in Figure (cid:6)(cid:4). For the bending bars printed with the organic binder system, a roller coater, as described in “Materials and Methods”, was used.Figure (cid:8)3: Organically in x-direction (see Figure 3) -D-printed bending bar after the --point-bending test. The protruding glass fibers are clearly visible, as well as the fracture.

The protruding glass fibers are clearly visible, as well as the fracture. The bending bar is still held together due to the presence of glass fibers.DiscussionSpecimen ProductionFor the selection of the right reinforcing fibers for the mold production, different glass fiber types were investigated regarding their reinforcing effect. The results are shown in detail in Table -. It was found that GF (cid:4) with (cid:4). mm fiber length (uncoated glass fiber) was the best fiber regarding high bending strength, but a big drawback was the huge agglomeration of the uncoated fibers. A conducted pull-out test supports the theory, that the uncoated fibers are good as reinforcing fibers for inorganic binder systems since there is no sizing on the fibers which might lead to a poorer fiber-binder adhesion. The uncoated fibers are unfortunately not well miscible with sand and binder.

The uncoated fibers are unfortunately not well miscible with sand and binder. This is the reason why fibers like GF were investigated, too, here the mixing was easier and a homogenous mixing was possible. It was found that this type leads to a much higher elongation at break at Fmax if enough fibers were used in the sand fiber mixture. The needed minimum amount of glass fiber for a higher elongation at break was .(cid:2) wt%, which can be seen in Figure 7. The use of a higher amount of binder did not lead to a higher elongation at break, which can be seen in Figure <. Of all the fiber types studied, GF2 was the most suitable for mixing sand, fiber and binder, and was therefore studied in depth, and the required quantities of1612this fiber for high flexural stress were researched in detail. These results were then used for mold production.

These results were then used for mold production. In theory, recycled glass fibers (GF 7) could be used for this application, provided that it's feasible to mix these fibers and the recycled fibers are free from impurities on the surface. Milled glass fibers (GF (cid:2)) are not suitable, since they are too short to show the effect of creating a quasi-ductile breaking behavior. GF (cid:6), GF (cid:28), and GF (cid:5) also showed agglomeration effects, so they were not further investigated. However, with a new mixing technology, those fibers might lead also to good results regarding the improvement in a higher elongation at Fmax.Casting of Molds and Surface DeviationsIn Figure 7, we can see that the elongation at break increases with a higher proportion of glass fibers.

Based on the casting experiments, we observed that a fiber-reinforced mold with a fiber content of (cid:28) wt.% could only be used for four castings, while the fiber-reinforced molds with (cid:2) wt.% and (cid:3) wt.% could be used for at least ten castings. This suggests that the elongation at break has a decisive influence on the moldability of fiber- reinforced molds. The bending bars produced parallel to the casting molds do not show this strong dependency of elongation at break on the fiber content (Figure (cid:8)C). This may be attributed to the manufacturing method of the bending bars. However, it can be claimed that fibers in the mold material enable multiple castings.It was also observed that the fiber-reinforced molds exhibit increasing deviation from their original shape with each casting. This is likely due to the heat causing the water glass bonds in the top layer, adjacent to the melt, to decompose and release water, making the material system more brittle.

This is likely due to the heat causing the water glass bonds in the top layer, adjacent to the melt, to decompose and release water, making the material system more brittle. This would support the hypothesis put forward by C. Appelt [ !]. As a result, the top layer crumbles after each casting, leading to more erosion of the mold with each casting process.This effect can also be observed in the images generated using GOM ATOS Compact Scan and the CloudCompare software, which can be seen in Figure (cid:8),. It can be clearly noticed that especially at sharp edges and corners, where the heat input of the melt is the highest, the hollowing phenomenon is also the most prominent. For example, in Figure (cid:8),, it can be seen that the casting flash significantly increases with each casting.

For example, in Figure (cid:8),, it can be seen that the casting flash significantly increases with each casting. Additionally, the details, such as the divisions of the mane of the knight, which are clearly visible in the first casting, become hardly noticeable in the tenth casting.One possible way to slow down this process could potentially be the use of coatings that act as heat insulators between the melt and the fiber-reinforced mold. Furthermore, it is anticipated that a refractory coating would significantly improve the surface quality of the casting. The deviations from the ideal form could be reduced through centering. To better investigate the effects of thermal expansion on the mold, a more sophisticated model would be necessary.

The deviations from the ideal form could be reduced through centering. To better investigate the effects of thermal expansion on the mold, a more sophisticated model would be necessary. This needs to be further investigated in the next step.Binder-Jetting of Fiber-reinforced Bending BarsIt could be demonstrated that GF fibers can be processed in the (cid:28)D printing process using both binder systems (water glass and furan resin) and two different coating systems (roller and blade) with a layer thickness of (cid:28)(cid:2)! µm or more.It is shown that with the (cid:28)D printing of fiber-reinforced bending bars using an organic binder system, a higher average elongation at break could be achieved compared to the reference bending bars without fiber reinforcement.

It is also evident that the fiber-reinforced bending bars exhibit a quasi-ductile stress-strain behavior, which can be seen in the left diagram of Figure (cid:8):.It is unclear why the (cid:28)D printed bending bars, printed with the inorganic binder, exhibit brittle fracture behavior, as seen in the right diagram in Figure (cid:8):. This could be due to the use of a different coating module used by the prototype in-house production (cid:28)D printer compared to the VX(cid:6)!!! (cid:28)D printing system. The prototype (cid:28)D printer for the organic binder system spreads the sand onto the underlying sand layer that has already been sprinkled with glass fibers. It was observed that some of the fibers were still in a vertical position – possibly due to electrostatic forces – before the sand coating process. During the sand-spreading process, these vertical fibers were able to maintain their position and form a connection between the layers.

During the sand-spreading process, these vertical fibers were able to maintain their position and form a connection between the layers. By establishing fiber connections between the individual sand layers, the fibers can create a higher resistance to cracking, leading to a quasi-ductile material behavior. This could be a possible explanation for the quasi-ductile stress-strain behavior, which can be seen in the left diagram in Figure (cid:8):. Another possibility is that a too small quantity of fibers was introduced1613through the manual process using a sieve. This would also result in the stress-strain curve shifting towards brittle failure.

This would also result in the stress-strain curve shifting towards brittle failure. In further experiments, it is advisable to determine the quantity of fibers introduced to rule out this scenario.organic binder systeminorganic binder system54fiber reinforced reference54reference fiber reinforced] a P M3] a P M3[ σ2[ σ2110 0,0%0,5%1,0%1,5%0 0,0%0,5%1,0%ɛ [%]ɛ [%]Figure (cid:8):: The left diagram shows the comparison of two representative stress-strain curves between a fiber-reinforced bending bar and a reference bending bar. The bending bars were -D printed using an organic system. The right diagram displays the comparison of two representative stress-strain curves between a fiber-reinforced bending bar and a reference bending bar.

The right diagram displays the comparison of two representative stress-strain curves between a fiber-reinforced bending bar and a reference bending bar. The bending bars were -D printed using an inorganic binder system.In contrast, the stress-strain curves of the bending bars produced with an inorganic binder system showed a brittle behavior, regardless of whether they were fiber-reinforced or not. This can be seen in the right diagram of Figure (cid:8):. In this case, the fiber-reinforced bending bars also exhibited lower elongation compared to the reference bending bars. The average elongation at maximum bending stress was !.(cid:6)(cid:2) % for the fiber-reinforced bending bars and !. ! % for the reference bending bars. This was not expected. Again, the coating process could provide a possible answer. In the VX(cid:6)!!! (cid:28)D printing system, the coating module dispenses sand from the front, which is then smoothed out with a blade.

Again, the coating process could provide a possible answer. In the VX(cid:6)!!! (cid:28)D printing system, the coating module dispenses sand from the front, which is then smoothed out with a blade. This sand wall created in front of the coating module smooths out all the fibers, including those that were in a vertical position. As a result, a pure layer of fibers is formed, sandwiched between two layers of sand. This does not allow for a connection between the sand layers through the fibers. The formation of these separate layers, sand-fiber-sand, prevents the fibers from providing resistance during crack formation. The sand layers fracture in a brittle manner, resulting in the failure of the entire bending bar.It should be mentioned that three bending bars were produced with the inorganic binder system, while only two bending bars were produced with the organic binder system. Therefore, only a trend can be observed in this case as well.

Therefore, only a trend can be observed in this case as well. In order to statistically validate these results, further experiments are urgently needed. Furthermore, the glass fibers were manually sieved, so the exact amount of fibers per bending bar cannot be determined, and the amount of fibers introduced can vary greatly between the produced bending bars, which can also result in significant variation in the results of the (cid:28)-point bending test. This was also observed in the hand- formed bending bars. An amount of at least (cid:28).(cid:3)(cid:2) wt.% glass fibers was necessary to observe a quasi-ductile material behavior in the fiber-reinforced bending bars.SummaryDue to the continuous increase in sand consumption, a finite resource, it is desirable to guarantee resource-saving through further processing and reprocessing.

To create an environmentally friendly and sustainable production chain in sand casting, new solutions for the binder systems and an efficient use of resources through reusable casting should be found. The aim of reinforcing sand molds and cores with fibers is to find a middle ground between the immense investment for permanent molds and the single use of sand molds.It has been shown that the use of glass fibers generally reduces bending strength but increases elongation at break. This effect can be enhanced by increasing the addition of fibers. In summary, coated glass fibers (GF )16141,5%with a length of (cid:5).(cid:28) mm yielded the best results. Furthermore, it has been observed that a larger quantity of binder increases bending strength, but does not significantly alter elongation at break.By adding glass fibers and thereby altering the material properties of the sand-binder system, bending beams with a quasi-ductile material behavior could be produced.

This quasi-ductile material system is suitable for the production of sand molds for multiple castings, as demonstrated. Fiber-reinforced sand molds with a fiber content of (cid:2) wt.% and (cid:3) wt.% could withstand at least ten castings. It was observed that the mold was progressively damaged with each casting, but it did not completely break until the tenth casting. In contrast, reference molds consisting only of sand and binder without fiber content were completely destroyed after the first casting. There was a tendency that a higher fiber content allowed for a higher number of castings per mold, but this result still needs to be statistically validated.The use of glass fibers was also demonstrated in the sand-based binder jetting process. Two differentbinder systems and coating processes were tested. It was shown that also in 3D printing the use of short fibers in combination with an organic binder system leads to a quasi-ductile behavior of the produced bending beams.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFIELDABLE PLATFORM FOR LARGE-SCALE DEPOSITION OF CONCRETE STRUCTURESPhillip C. Chessera, Brian K. Posta, Alex Roschlia, Randall F. Linda, Alex M. Boulgera, Lonnie J. Lovea, and Katherine T. GaulaaManufacturing Demonstration Facility, Oak Ridge National Laboratory, 2370 Cherahala Blvd. Knoxville, TN 37932AbstractOak Ridge National Laboratory’s Manufacturing Demonstration Facility is developing a novel, large-scale additive manufacturing, or 3D printing, system. The Sky Big Area Additive Manufacturing (SkyBAAM) system will ultimately be a fieldable concrete deposition machine with pick and place abilities that will allow for full-scale, automated construction of buildings. The system will be implemented with existing construction equipment meaning conventional cranes will be used to suspend the print head.

The system will be implemented with existing construction equipment meaning conventional cranes will be used to suspend the print head. SkyBAAM will be cable-driven by four base stations and suspended from a single crane. The elimination of a gantry system, found commonly in large-scale additive manufacturing systems, will enable SkyBAAM to be quickly set up with minimal site preparation. The medium-scale version of SkyBAAM is currently in development. The system design, cable stiffness analysis, and tactics for freezing rotational degrees-of-freedom (DOF), detailed in this paper, will provide a basis for the final, large-scale version of the SkyBAAM system.IntroductionAdditive manufacturing (AM) has the potential to revolutionize and revitalize American manufacturing. Most conventional methods of AM are currently only practical to produce small, high-value, and low-volume components. This leads to applications such as aerospace and biomedical.

Most conventional methods of AM are currently only practical to produce small, high-value, and low-volume components. This leads to applications such as aerospace and biomedical. More recently, AM has been used to fabricate tooling for a broad range of applications. [1] This includes tooling for automotive and other high-volume applications.However, this scope excludes one of the largest sectors of the U.S. economy, namely the construction industry. New construction in the U.S. is projected to grow to $1.4 trillion by the year 2021. [2] This is roughly 6% of the U.S.’s gross domestic product (GDP). AM fabrication of structures or buildings has the potential to revolutionize this market, saving labor and material cost and shortening construction time.The idea of AM for construction has been around for decades; although, it is currently far from becoming a real-world application.

In the framework presented below, structures are deposited by an extruder, one layer at a time, with a highly-viscous, low-slump concrete. This eliminates the need for costly formwork. Ideally, as the building is deposited, pick and place systems will lay components for the plumbing and electrical systems of the building in channelsThis manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes.

The Department of Energy will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).2020that have been left for this purpose. Rebar and other reinforcing components can also be laid in with pick and place systems. Without the need for formwork, new geometries, impossible with conventional methods, can be realized. This makes new, energy-efficient structures possible.Construction waste reduction is also a possibility with AM. Waste from new construction in the U.S. in 2014 was 28.9 million tons. [3] With AM, material is only placed where it is needed; therefore, AM is inherently low waste. There is no need for traditional cutting or machining of stock parts and throwing out the waste. There is also no need for temporary structures, such as formwork, that are later discarded.

There is no need for traditional cutting or machining of stock parts and throwing out the waste. There is also no need for temporary structures, such as formwork, that are later discarded. Large-scale penetration of AM into the construction industry has the potential to greatly reduce construction waste.The end vision for construction-scale AM is grand. It is easy to envision energy efficient [4], aesthetically pleasing buildings constructed with less waste using automated processes that reduce labor. The potential effect on the economy could be profound. Take the following example. Currently, houses cost more than the average yearly income of the residents who buy them. Yet automobiles, which are massively more complex than houses, cost only around a tenth of what housing costs. Part of the reason for this is that automobiles are produced with the benefits of large-scale automation. This is not true with construction, which is a manual, labor intensive industry.

Part of the reason for this is that automobiles are produced with the benefits of large-scale automation. This is not true with construction, which is a manual, labor intensive industry. The application of automation to construction through AM has the potential to make housing drastically more affordable. Similar benefits could be seen with commercial construction, where lower building costs could lead to increased ease of expansion of businesses and the creation of jobs. Unfortunately, the realization of this vision is still a long way off and requires much more research.Research on concrete AM started as early as the 1990s; although, the idea was not widely accepted. Randall Lind, a researcher at Oak Ridge National Laboratory (ORNL), first recorded the idea in 1993, and presented work in this field in 2009.

Randall Lind, a researcher at Oak Ridge National Laboratory (ORNL), first recorded the idea in 1993, and presented work in this field in 2009. [5] Several years after Lind’s initial conception of the idea, Khoshnevis [6] started work on what is now known as Contour Crafting.In recent years, concrete AM has been gaining traction, and there have been continued advances in this field, including advances in materials and extrusion techniques. However, concrete AM is still mostly explored and used in a laboratory setting and has not been seen in industry. Unfortunately, the scale necessary for effective use in construction remains elusive. A large reason for this is because motion platforms used for deposition are highly inadequate for the needs of the construction industry, as will be explored below.

A large reason for this is because motion platforms used for deposition are highly inadequate for the needs of the construction industry, as will be explored below. This paper aims to examine the possibility of using cable-driven robots for deposition of concrete in the construction industry as a means of making concrete AM a commercially feasible process.State of the Art: Problems and AlternativesAs mentioned above, work on concrete AM for construction is not new. In fact, there is much prior work in the field. This will be briefly surveyed. However, there is a significant lack of solutions that are practical for use on a real-world jobsite.2021Current Additive Manufacturing of Concrete StructuresAM of concrete for construction has been a popular area of research in recent years because this process promises to benefit the construction industry in many ways. Biernacki et. al.

Biernacki et. al. have explored some of these possible benefits, such as safety, efficiency, and the ability to fabricate new types of structures. [7] Khoshnevis and the Contour Crafting group are old players in the field. [8] They have been successfully 3D printing with concrete for many years and have proposed many applications for the technology. The U.S. Army Corps of Engineers is researching concrete AM as well, with an interest in printing barracks and forward operating bases. [9] Another player in this area is WinSun, a Chinese company that is perusing the commercialization of concrete AM. Their work includes the printing of an office building in Dubai among other projects. The growing number of groups working in the field of concrete AM has been noted by Bos et. al.

Their work includes the printing of an office building in Dubai among other projects. The growing number of groups working in the field of concrete AM has been noted by Bos et. al. [10]a)b)c)Figure 1: a) US - Army Corp of Eng., b) Russia - Apis Cor, c) China – WinsunMany of the underlying fundamental material and extrusion problems associated with AM of concrete have also been examined. For example, work has been done to study the mechanical properties of additively manufactured cement [11], [12] as well as the printability of different concrete mixes. [13]The success of many of the players in the concrete AM realm led the authors to conclude that AM of cement structures is a problem that is well on its way to being solved. The big question that remains pertains to the practicality of 3D printing concrete structures.

The big question that remains pertains to the practicality of 3D printing concrete structures. Can concrete AM be effectively implemented on the jobsite? This paper aims to answer that question in the affirmative by presenting a concept for a fieldable deposition platform.The Impracticality of Gantry Based Deposition for ConstructionWhile the deposition of concrete for AM structures has been effective, concerns arise when considering the current motion platforms within this context. Nearly all material extrusion (ME) AM systems use a gantry-based motion platform, and large-scale systems for deposition of structures are no exception.Scaling gantry systems to a sufficient size to manufacture buildings while making the systems fieldable creates problems. Gantry systems, when produced on this scale, become expensive. Furthermore, fielding gantry systems is quite difficult. Since the structure is printed within the gantry itself, it must be larger than the structure.

Furthermore, fielding gantry systems is quite difficult. Since the structure is printed within the gantry itself, it must be larger than the structure. For large buildings, this requires very large machinery that must be transported to the jobsite and installed, likely at a very high cost. The time required to set up a gantry on the jobsite would also likely be significant. Once on the jobsite, the gantry would have to be placed on level ground. This limits the terrain a2022gantry could be deployed on. Even on relatively level ground, there would be an added cost to get the site level within the accuracy required for a gantry system to operate. Overall, there are significant challenges and costs associated with deploying a gantry system to a jobsite.Cable-Driven RobotsIt is desirable to have a fieldable AM system that can be deployed with ease on a jobsite, allowing a structure to be printed on site.

Gantry robots are unlikely to provide this at an economical cost for reasons discussed above. Thus, it is desirable to find another type of motion platform that can be used for this application.Apis Cor has proposed an alternative to a Cartesian gantry-based system by using a polar style robot. [14] This approach has also been used by Neri Oxman’s team at MIT. [15] While this eliminates some of the disadvantages of a traditional gantry system, it has fundamental size and stiffness limitations. A good motion system for deposition is still needed.An alternative solution can be found with cable-driven robots. Cable-driven robots or cable-driven parallel manipulators (CDPMs) are becoming increasingly popular. They provide potentially large workspaces with relatively light equipment. [16] By using cables, workspaces much larger than the machinery itself can be achieved.

They provide potentially large workspaces with relatively light equipment. [16] By using cables, workspaces much larger than the machinery itself can be achieved. This is different from gantry robots, where the machinery must be larger than the workspace.Cable-driven robots have been used in numerous applications. One notable example is the Skycam, which uses cables to suspend a camera over a sports arena and moves the camera to follow play. [17] Full six degree-of-freedom (DOF) robots have also been developed. [18] A notable example is the National Institute of Standards and Technology (NIST) robocrane, [19] which has spawned large area variants. [20] CDPMs have also been used in large telescopes, which further demonstrate the large-scale at which they can be used. [21]Cable-driven robots are a potentially suitable replacement for very large-scale AM applications for several reasons.

[21]Cable-driven robots are a potentially suitable replacement for very large-scale AM applications for several reasons. Most importantly, they easily achieve large workspaces without requiring massive equipment and machinery. Additionally, cable-driven robots do not require the same level of site preparation as gantry robots. With a CPDM based system, cable winders can be placed at several base stations around the worksite. This would not require leveling of the site and could be implemented on a wider variety of terrain. Furthermore, these base stations would be small in comparison with the workspace, leading to a lower machinery cost for a given build volume.

Furthermore, these base stations would be small in comparison with the workspace, leading to a lower machinery cost for a given build volume. In the remainder of this paper, a scheme for a fieldable, cable- driven robot for large-scale AM is examined.SkyBAAM Deposition PlatformSky Big Area Additive Manufacturing (SkyBAAM) is a proposed cable-driven motion platform that is designed specifically around the requirements for the deposition of large-scale AM structures. These specific objectives will the explained below. Large-scale AM with cable- driven platforms is not a new idea. [22] Bruckmann et. al have also proposed laying bricks with cable-based systems [23], while Vukorep has realized a cable-driven system that places foam blocks. [24] However, the SkyBAAM system proposes a radically different cable architecture from these systems. The SkyBAAM cable system is designed to be easily fieldable at a jobsite with as much conventional equipment as possible.

The SkyBAAM cable system is designed to be easily fieldable at a jobsite with as much conventional equipment as possible. The cable architecture was inspired by the Hangprinter created by Torbjørn Ludvigsen. [25] However, the requirements of the SkyBAAM2023system are different than the Hangprinter, so there have been significant departures from the original Hangprinter.SkyBAAM ObjectiveThe overall objective of SkyBAAM is to create a cable-driven motion platform that is easily fieldable for the additive manufacturing of large structures. Furthermore, it is also designed to integrate with existing construction equipment. An initial conceptual design is shown in Figure 2.Figure 2: SkyBAAM conceptSingle Lift PointMost cable-driven robots have cable winders at multiple aerial points. In a situation where the robot is permanently installed in a building or high-bay, this is not a problem as these winders can be affixed to a large frame or to the building it is housed by.

In a situation where the robot is permanently installed in a building or high-bay, this is not a problem as these winders can be affixed to a large frame or to the building it is housed by. However, in an outdoor fieldable platform this is less practical. Bruckman et. al. [23] proposed the erection of a large frame around the area where the structure will be fabricated to provide cable attachment points. However, this approach would require significant time and labor to erect a frame that is both large and stiff enough. It is desirable to eliminate the need for such a structure, which increases the ease of fieldability and reduces the amount of time required to set up the system on site.The best solution is to have a single aerial winding point or platform that can be suspended by a mobile crane. To integrate this system with existing construction equipment, the crane must not be part of the motion system.

To integrate this system with existing construction equipment, the crane must not be part of the motion system. Instead, it simply acts as a static hoist point for a winder, which is integrated into the motion control. Because the crane merely provides a hoist point for the system, it allows the SkyBAAM system to be integrated with any crane that meets the necessary height and load requirements.It is also desirable to keep the total number of winders and cable anchor points to a minimum to reduce the number of pieces of equipment required. Furthermore, by concentrating cable winding points and anchor points together, the system can be split into a handful of base stations for deployment in the field. These base stations can be integrated into a truck or trailer that can be driven onto the site.2024Constraint and Degrees-of-FreedomAny object in 3D space has six DOF. To fully constrain the end effector, most cable- driven systems control all six DOF.

To fully constrain the end effector, most cable- driven systems control all six DOF. However, this requires six independently controlled winders. For most AM, only three DOF control is needed. Linear translation is necessary in all three directions, however, the three rotational DOF are not needed. Instead of controlling the rotational DOF, it is more appropriate to fix them and only control the linear translation. This can be achieved, by carefully choosing the cable geometry and controlling multiple cables from one winder. By only controlling the three linear DOF, it is possible to make a simpler system. This, in turn, aids in the ease of fieldability.Over-constraint is also a problem that must be avoided. If there are more cables than necessary to fully constrain the system, it becomes over-constrained. In an over-constrained system, forces can easily get larger than would otherwise be the case, and the demands on the control system are much greater.

In an over-constrained system, forces can easily get larger than would otherwise be the case, and the demands on the control system are much greater. Thus, it is desirable to have an exactly constrained system.Cable ConfigurationThe cable configuration was chosen to meet the constraints already mentioned. Insummary these are:1. Single aerial winder that can be held by a standard crane 2. Concentrate cable winders and anchor points into several base stations 3. Fully and exactly constrain the end-effector 4. Freeze all rotational DOFWhile meeting these requirements, the stiffness of the system must be kept high to maintain good print quality. The importance of high stiffness is addressed in further detail below.Figure 3 shows the initial cable configuration. The red body represents the print head. Three vertical “z-cables” suspend the print head and control the vertical motion. These three cables all spooled on one winder to ensure their lengths are equal.

Three vertical “z-cables” suspend the print head and control the vertical motion. These three cables all spooled on one winder to ensure their lengths are equal. This keeps the print head from tilting, which fixes two rotational DOF. This winder is located on the yellow platform, which is held up by the crane. Six stay cables, shown in blue, prevent this platform from moving.Figure 3: Initial cable configuration for SkyBAAM2025Motion in the horizontal plane is controlled by the three cables wound on the blue motion control stations. One of these stations has two equal length cables. This fixes the third rotational DOF.A cable from the orange station provides tension only. The purpose of this tension cable is to provide a nesting force that keeps all the cables under tension. It does not contribute to the motion. Thus, this extra cable does not cause the system to be over-constrained.Figure 4 shows the final cable configuration.

It does not contribute to the motion. Thus, this extra cable does not cause the system to be over-constrained.Figure 4 shows the final cable configuration. This configuration has several advantages incomparison to the previous configuration.Figure 4: Final cable configuration for SkyBAAMThe two extra cables in the vertical direction in the initial configuration are brought down to the horizontal plane in this configuration. This still freezes all the rotational DOF but offers several more advantages. First, by putting more of the cables in the horizontal plane, the stiffness is increased in the horizontal plane. This does sacrifice some of the vertical stiffness, but the horizontal stiffness is more important in this system because most of the movement will be in the horizontal plane. The only vertical movement will be between layers when the print head is raised by a small increment after each layer.This change also simplifies the system by eliminating the upper platform.

The only vertical movement will be between layers when the print head is raised by a small increment after each layer.This change also simplifies the system by eliminating the upper platform. With only one vertical cable, it is possible to run this cable over a pulley and down to a winder located on a ground base station. This pulley is still held up by the crane, but it is possible to hold it stationary with only three stay cables rather than six.The final change was to add a second cable to provide tension. Both the tension stations are shown in orange in the figure. The net force required to keep the other cables in tension is the vector sum of the tension in these two cables. By adjusting the tension of the two cables, the net tension vector can by adjusted to some degree to optimize the system. The importance of this will be discussed in a later section.Layout and Deployment in the FieldAs noted before, an objective of this system is to be easily fieldable.

The importance of this will be discussed in a later section.Layout and Deployment in the FieldAs noted before, an objective of this system is to be easily fieldable. The four base2026stations could be contained on individual trailers. These trailers could drive into the correct positions on site. The overhead point can be suspended by a crane. After this, the cables would be attached to the print head, and the machine would go through a calibration routine. Then it would be ready to print.Material must be delivered to the head during deposition. A simple way to achieve this is to have a hopper on board the print head. This hopper will be periodically refilled as it uses material. This introduces some control challenges because the weight of the print head will be constantly changing. A better possible method would be to continuously supply material to the print head as deposition is taking place. This could be done with a concrete pumping truck.

A better possible method would be to continuously supply material to the print head as deposition is taking place. This could be done with a concrete pumping truck. The boom of the pumping truck would extend above the system. A flexible hose would then run all the way from the extended boom to the print head and provide a continuous flow of material for deposition.Design ChallengesSo far, this paper has explained the objective and the general architecture of the SkyBAAM. A scaled-down, operational system is currently in development at ORNL’s Manufacturing Demonstration Facility. The remainder of this paper will discuss some of the design challenges that must be addressed to make SkyBAAM a reality. The purpose is not to present a fully-developed design but to examine some of the important challenges and how they are being solved.StiffnessA major concern in the design of SkyBAAM is the effective stiffness of the end-effector.

The stiffness will affect both the quality of prints and the speed at which deposition can occur. In a stiffer system, the end-effector will deflect less under dynamic loads, leading to a more accurate and repeatable system. Furthermore, a stiffer system leads to higher natural frequencies, which makes control of the system easier.As mentioned in a previous section, the cable layout of the system was chosen specifically to increase the stiffness of the system in the x-y plane, as this is the plane in which most of the motion occurs.Stiffness is a driving factor in the design of the whole SkyBAAM system. Not only was the cable configuration selected based on stiffness requirements, but throughout the whole design process, stiffness was a driving consideration.Cable Selection and Catenary EffectsCable selection is important for the SkyBAAM because high stiffness is desirable for this system. Cable selection and cable tension play significantly into the stiffness of the system.

Cable selection and cable tension play significantly into the stiffness of the system. In selecting a cable and the required tension, the effects of catenary sag must be considered.A cable span under gravity sags in a well-known and well-studied shape called the catenary. [26] The amount of sag is dependent upon the amount of tension in the cable as well2027as its weight and length. Catenary effects result in significant non-linearities in the stiffness of cables. [27]The authors combined the non-linear stiffness produced by catenary sag with the elastic stiffness of the cable to produce a composite stiffness value that is dependent upon cable tension. The mathematical derivation for this will be detailed in a following publication. An example of this for a 17ft long segment of 1/8-inch 7x19 steel wire is plotted in Figure 5.) n i / s b l (s s e n f f i t STension (lbs)Figure 5.

An example of this for a 17ft long segment of 1/8-inch 7x19 steel wire is plotted in Figure 5.) n i / s b l (s s e n f f i t STension (lbs)Figure 5. Stiffness of 1/8” 7x19 steel wire rope 17’ long varying with tensionAs shown in Figure 5, there is a point where the contribution of the sag to the stiffness becomes negligible, and the stiffness remains essentially constant with increasing tension. This observation is consistent across different cable types and sizes, while the exact shape of the graph changes. This makes it possible to classify cable stiffness into a non-linear range and a linear range.It is desirable to operate in the linear range for several reasons. First, this is the highest stiffness range of the cable. As stiffness is a priority, this alone provides significant motivation to be in this region. Furthermore, analysis and control of linear systems is significantly easier when compared to non-linear systems.

Furthermore, analysis and control of linear systems is significantly easier when compared to non-linear systems. For these reasons, it was decided that the cables of the SkyBAAM should operate in the linear region.The amount of tension required to reach the linear stiffness region of a cable as well as the value of the stiffness within the linear region are determined by the cable itself. The elastic stiffness, weight, and length of the cable play a significant role in the amount of tension required. To complicate things further, the elastic stiffness of the cable is affected not only by the material, but also by the way the fibers are woven within the cable. The interplay of these factors and their effect on the overall system make the selection of cable an important factor in the system design.2028Cable Tensioning SchemeAs mentioned in the previous section, it is important that cables operate within the linear range of stiffness.

To achieve this, all the cables must be kept at a certain minimum tension for a given cable type and length.There are two tensioning cables that keep all the cables in tension. As noted above, this allows for some control of the net tension vector. Tension control in cable-driven robots has been proposed by others. [28] [29] Here, we use the control of tension to keep all the cables above the minimum tension required for operation in their linear stiffness region. However, there is more than one solution that meets this requirement. The set of all possible solutions is found in the null space of the static equilibrium matrix. This method will be detailed in a later paper. Some of the solutions may involve tensions in a few of the cables that are well above the minimum tension. However, it is not desirable to tension cables more than required. This produces no benefit; over-tensioning cables only serves to increase the required motor power and the required strength of the components.

This produces no benefit; over-tensioning cables only serves to increase the required motor power and the required strength of the components. Thus, the solution that keeps all the cables above their minimum required tension and has the lowest maximum tension is chosen. The required force in the tensioner cables will vary as the end-effector traverses the workspace.Motion TrackingAnother important consideration in the design of the SkyBAAM is how to track the location of the end-effector through 3D space. It is also imperative to know the location of the winders to solve the inverse kinematics. Both issues will be addressed with a laser-based time- of-flight sensor. A Leica laser tracker will be used to locate and track the end-effector during deposition. It will also be used to locate the winders and base stations for initial calibration of the system.In addition to the use of laser-based time-of-flight sensors, the use of relative GPS will be examined.

Relative GPS can be used to obtain accurate relative positions but do not require line of sight like a laser-based system would.System Fabrication & Next StepsCurrently, a detailed design is being completed for a mid-scale SkyBAAM system at ORNL, and fabrication of that system has started. Assembly and testing are currently ongoing over the summer and fall of 2018. There are also plans to design and fabricate a full-scale prototype system with industrial partners that will then be tested in real-world scenarios.ConclusionThis paper presents the concept for a large-scale, cable-driven system for the deposition of structures from cementitious material. While recognizing the existence of prior art in large- scale AM for construction using cable-driven platforms, this system is designed to be easily fieldable for use on a jobsite.

While recognizing the existence of prior art in large- scale AM for construction using cable-driven platforms, this system is designed to be easily fieldable for use on a jobsite. A proposed cable configuration is chosen to achieve this objective, and associated design challenges are examined.2029This system is aimed at overcoming one of the major challenges in the AM of structures, namely the current lack of a fieldable motion platform. The development of a practically fieldable platform promises to be another step toward the goal of automated fabrication of structures with AM methods.The SkyBAAM system has been designed around needs specific to AM in the construction industry. This paper presents the overall design and layout of the SkyBAAM system. Currently, a mid-scale prototype is being designed and built at ORNL.

This paper presents the overall design and layout of the SkyBAAM system. Currently, a mid-scale prototype is being designed and built at ORNL. This prototype is the next step toward a full-sized system, which if successful, will be an important step toward realizing practical additive manufacturing of structures.AcknowledgementResearch sponsored by the Laboratory Directed Research and Development Program of Oak Ridge National Laboratory, managed by UT-Battelle, LLC, for the U. S. Department of Energy.2030References[1][2][3][4][5][6][7][8][9]Kunc V. el al. Large-Scale Additively Manufactures Tooling for Composites. Proceedings of 15th Japan International SAMPE Symposium and Exhibition. 2017. Forecast for new construction put in place in the U.S. from 2011 to 2021 (in billion U.S. dollars) (Rep.). (2017, July).

lament Feed Materials for Fused Deposition Processing of Ceramics and MetalsM.K. Agarwala, R. van Weeren, A. Bandyopadhyay, A. Safari, and S.C. Danforth Center for Ceramic Research, Rutgers UniversityW.R. PriedemanStratasysInc.Fused Deposition of Ceramics (FDC) and Metals (FDMet) are SFF techniques, based on commercial FDMTM technology, for fabrication of ceramic and metal components. The FD processes use feed material in the form of filaments which require certain phy~ical and mechanical properties. FDC and FDMet processes employ fila.n::ents formed from ?erannc ~r metal powders mixed with thermoplastic polymers. The thermoplastic polymers act as bmder dunng the FDC and FDMet processing in forming a green part. Development of green ceramic or metal fIlaments for FDC or FDMet processing involves three critical steps : selection of an appropriate binder chemistry, appropriate mixing procedures and fIlament fabrication techniques.

This study describes the properties required for filaments for successful FD processing and the approach taken in the development of a series of binder which meets these requisite properties for a wide range of ceramics and metals. Appropriate mixing and filament forming techniques are also discussed.I.IntroductionMost of the SFF techniques for fabrication of ceramic and metal parts employ polymeric binder systems to bond the ceramic or metal particles together in forming a green part [1]. The green part is then processed to remove the binder from the part and then sintered or infIltrated with a lower melting second phase [1]. However, fabrication of green ceramic or metal parts by various SFF techniques have been far from trivial. First, effort is required in developing the proper binder chemistry to allow SFF processing. In addition, effort may be needed in developing suitable dispersants for the specific particulate system and the binder chemistry.

In addition, effort may be needed in developing suitable dispersants for the specific particulate system and the binder chemistry. Once an appropriate binder chemistry and dispersant have been selected, the ceramic or metal powder systems are processed with the binder and dispersant to form the feedstock in a suitable physical form (powder, sheets, filaments) for specific SFF processing [1].Each SFF technique requires feedstock materials with certain physical and mechanical properties. .Poor selection of binder chemistry, dispersant, and poor processing of the feedstock can lead to unsuccessful SFF processing. This study describes the development of green ceramic and metal fIlament feedstocks for successful processing by Fused Deposition of Ceramics (FDC) and Metals (FDMet).

This study describes the development of green ceramic and metal fIlament feedstocks for successful processing by Fused Deposition of Ceramics (FDC) and Metals (FDMet). FDC and FDMet are SFF techniques, based on commercial FDMTM technology, which are being developed for a wide range of ceramics and metals to produce complex shaped ceramic and metal parts with functional properties [2-5]. This study discusses the various physical and mechanical properties required in the filament feedstock for successful FD processing. The study also describes the development of a binder system and the filament fabrication procedures employed for various ceramic and metal systems.ll.Filament Feedstock Characteristics for FD ProcessingThe commercial FDMTM systems fabricate wax and polymer parts using continuous spooled fIlaments as the feed material. The filament is fed through a position-controlled heated liquefier head which deposits molten material in precise locations [6,7].

The filament is fed through a position-controlled heated liquefier head which deposits molten material in precise locations [6,7]. For the filament to be successfully fed into the liquefier and extruded in molten form out of a nozzle, the filament must have certain properties. The fIlament, which is fed into the liquefier head by a pair of counter rotating rollers, not only acts as the feed material, but also serves as an efficient piston for extrusion. Therefore, for a certain material viscosity the filament must have sufficient stiffness and451strength for extrusion, or else the filament will buckle or break after passing through the rollers and not be successful in extruding material [5,7]. Hence, necessary, but not sufficient, conditions for successful FD processing are low viscosity and high stiffness.

Hence, necessary, but not sufficient, conditions for successful FD processing are low viscosity and high stiffness. In addition, the material must have good adhesive behavior for bonding between adjacent roads and layers of deposited material [5,7].Although filaments with low viscosity, high stiffness, and good adhesion characteristics for complete process optimization and are sufficient automation, the filament feedstock must also posses high flexural modulus and flexural strength [5,7]. High flexural modulus and strength are needed to enable continuous spooling of the filaments onto a spool to provide continuous feedstock supply for the FDMTM system.

High flexural modulus and strength are needed to enable continuous spooling of the filaments onto a spool to provide continuous feedstock supply for the FDMTM system. These properties are also necessary to allow continuous un-spooling of the filament from the spool for uninterrupted feeding (without breaking of the filaments) of the liquefier during fabrication.to establish FD processing feasibility,FDC and FDMet processing, using commercial FDMTM systems, require development of powderlbinder filament feedstocks with the above mentioned properties, namely, low viscosity, high stiffness, good adhesion behavior, high flexural modulus and strength. This is done through binder development and selection of dispersant. In addition to these constraints, development of binder(s) and dispersant for FDC or FDMet must also address issues commonly relevant to all green forming techniques.

In addition to these constraints, development of binder(s) and dispersant for FDC or FDMet must also address issues commonly relevant to all green forming techniques. The binder and dispersant have as high a solids loading as possible (>50 volume%), have good dispersion behavior, and must leave no residue after binder removal.III. Binders for FDC and FDMetThe binders for green forming techniques typically consist of many components; each added for a specific role in the final mix. In this study a four component binder system has been developed for FDC and FDMet processing. As shown in Table I, each of these components plays a specific role in the final formulation. Table I also shows the temperature range at which each of these components volatilize during binder removal. The amount of each component in the formulation used for FDC or FDMet is varied to tailor the properties needed to allow successful FD processing of specific ceramic or metal system.

The amount of each component in the formulation used for FDC or FDMet is varied to tailor the properties needed to allow successful FD processing of specific ceramic or metal system. The weight percent of each of these components varied in this study for FDC/FDMet is also shown in Table I. This binder system is designated RUx (x is the identification number of a particular binder formulation under consideration) [5].The particulate systems studied for FDC and FDMet processing to date include silicon Initial nitride, silica, lead zirconium titanate (PZT), stainless steel, and tungsten carbide-cobalt. feasibility studies for each of these particulate systems were done by using the RU1 binder formulation which contains 30% wax, 35% polymer, 15% tackifier, and 20% elastomer by weight. 50 to 65 volume % powders were mixed the RU1 binder and fabricated into filaments for FD feasibility studies. In addition, viscosity measurements were made on the mixed feedstock.

50 to 65 volume % powders were mixed the RU1 binder and fabricated into filaments for FD feasibility studies. In addition, viscosity measurements were made on the mixed feedstock. Based on the viscosity behavior and FD feasibility studies, the binder formulation were further optimized. Also, dispersants were selected and used for certain particulate systems to facilitate FD processing. Depending on the study done with each particulate system, the binder formulation development and dispersant selection was either limited to developing filaments which are FD process worthy but not flexible or was developed to produce FD process worthy and continuous and flexible filaments. Table IT shows the different RU binder formulations used successfully for FDC and FDMet processing of various materials.

Table IT shows the different RU binder formulations used successfully for FDC and FDMet processing of various materials. The Table also shows the amount of powder used for FD processing of each system and the nature of the filament feedstock thus produced.Binder and Dispersant Development for FDC of Silicon Nitrideto date has been done with Si3N4 [3-5]. Although initial FDC trials of 60 volume % Si3N4 with RU1 binder formulation were successful, the filaments were rigid and not flexible enough to enable spooling for continuous and automatedMost significant FD process development452FDC processing. To facilitate continuous and automated FDC processing of Si3N4, the RU binder was tailored and optimized by a linear Design of Experiments (DoE) [5]. Nine different RU fonnulations with a wide range of wax. polymer, tackifier, and elastomer contents were used.

Nine different RU fonnulations with a wide range of wax. polymer, tackifier, and elastomer contents were used. Viscosity measurements were made on each of these compositions as a function of te:n~erature and shear rate to get the viscosity response as a function of each of 0e co~ponen.ts. Slt~:nlarly, gI~ss transition temperatures (Tg) were measured for each fonnuiatlon uSIng D~ferentlal Scannmg Calorimetry. Tg is commonly used as an indicator of the flexi~ility of the matenal. The. d~?t from the DoE was analyzed to determine the effects of the vanous components on fleXIbility and viscosity. Using a linear representations of the response,Response =Co + Cl [tackifier] + C2 [elastomer] + C3 [polymer] + C4 [wax]the coefficients (Co -(4) for the effect of each component on viscosity and flexibility were calculated [5]. From this linear DoE study, the following conclusions were drawn:The tackifier and elastomer improve flexibility and the wax adversely affects the flexibility.

From this linear DoE study, the following conclusions were drawn:The tackifier and elastomer improve flexibility and the wax adversely affects the flexibility. • Elastomer increases the viscosity and the wax lowers the viscosity.Based on the results of DoE, the binder system RU9 was developed. RU9 formulation contains 20% wax, 19% polymer, 35% tackifier, and 26% elastomer by weight. In addition to developing RU9 binder formulation, a suitable dispersant was also selected. Selection of oleyl alcohol as dispersant was done by a screening process from among several commonly used dispersants for ceramics, especially silicon nitride [5]. The screening of dispersants was done by measuring the viscosity of 30 volume % Si3N4 loaded RU binder suspensions. For the dispersant screening experiments, the powder was pre-treated with varying amounts of different dispersants, prior to mixing in RU binder.

For the dispersant screening experiments, the powder was pre-treated with varying amounts of different dispersants, prior to mixing in RU binder. Based on these screening experiments by measuring viscosity, as shown in Figure 1, 3 weight % oleyl alcohol, which exhibits the greatest decrease in the viscosity, was detennined to be a suitable dispersant [5,8]. Hence, the Si3N4 powders used in the study were pre-treated with 3 weight % oleyl alcohol, prior to mixing with the RU9 binder formulation. As shown in Figure 2, filaments fabricated from green feedstocks, containing 45% RU9 binder and 55% Si3N4 powder by volume, pre-treated with 3 weight % oleyl alcohol, were flexible, continuous and easily wound onto a spool for successful FDC processing.

Although the viscosity of Si3N4 FDC formulations were considerably higher than that of an FDMTM material, investment casting wax, Figure 3, FDC processing of these materials was equally successful.Binder Development for FDC and FDMet of Other MaterialsAs indicated in Table IT, binders for other material systems (silica, PZT, stainless steel, and tungsten carbide-cobalt) have not been optimized to result in continuous and flexible filaments. The binders for these materials were developed with appropriate viscosity to allow successful FD processing using small pieces (6"-12" in length) of stiff and rigid filaments, Figure 4. FDC and FDMet processing of 65 volume % silica and 60 volume % stainless steel was feasible with the RU1 binder without the use of any dispersant. However, the RU1 binder was tailored and developed to allow FD processing of PZT and WC-Co.

However, the RU1 binder was tailored and developed to allow FD processing of PZT and WC-Co. Initial FD analysis of 50 to 55 volume % PZT and WC-Co with RUl binder indicated viscosity levels much higher than permissible for successful FD processing. Therefore, suitable dispersants have been selected for PZT and WC-Co to lower the viscosity of 50 volume % powder in RUl binder composition. Figure 5 shows such a lowering o~visc?sity of 50 volume % WC-Co when pre-treated with 1 weight % dispersant. Use of appropnate dIspersants for PZT and WC-Co have enabled successful FD processing with 50 volume% loading in RUl binder, although the filaments remain rigid. As was done for silicon nitride, it is expected that appropriate dispersant selection and RU binder optimization would result in continuous and flexible FD process worthy filaments for these materials and for other ceramic or metal systems.453IV.

Compounding of Ceramic or Metal Powders with BindersOnce the powder, binder and dispersants have been selected for the FDC/FDMet process, the next concern is homogeneous mixing of these ingredients. Mixing or compounding is crucial, since deficiencies in the quality of mixed feedstock can not be corrected by subsequent processing adjustments. Failure to deagglomerate and evenly distribute the powder at this stage will result in downline defects in both FDC fabrication and post-processing operations. Various techniques, such as, measurement of mixing torque or energy, shear modulus, viscosity, etc. are used to assess the quality of mix and to ensure consistent and repeatable compounding from batch to batch.Both heat and shear are required for efficient compounding of RU-type thermoplastic binders and the ceramic or metal powders. Therefore, torque rheometer mixing was selected as the method for compounding FDC/FDMet feedstocks.

Therefore, torque rheometer mixing was selected as the method for compounding FDC/FDMet feedstocks. Torque rheometer mixing also measures the torque of mixing which is used in assessing the quality of mix. When dry powder is first added to the molten binder, a high torque value is recorded due to entrapment of the binder between and in the powder agglomerates. As the agglomerates break down by the shearing action of the blades, the immobilized entrapped liquid binder is released and hence the torque goes down to a constant temperature of level and stabilizes. The stabilized torque is determined by the solids loading, mixing, speed of mixing, and the dispersing qualities of the binder.

The stabilized torque is determined by the solids loading, mixing, speed of mixing, and the dispersing qualities of the binder. Any change in mixer speed (shear rate) and temperature will cause the torque to change since the system viscosity, and hence the torque, are shear rate and temperature dependent.Experimental Procedure and Results for Compounding:In this study, the compounding procedure for incorporating ceramic or metal powder into the binder system involves the use of Haake Rheocord torque rheometer [9]. The torque rheometer is fitted with batch mixing bowls of capacities 50 cm3 and 250 cm3, with Z-blades to provide shear mixing. The maximum blade speed of the mixers is 150 rpm and the maximum temperature capability is 200 oe. The mixing bowls are double walled with three independently heated zones.

The maximum blade speed of the mixers is 150 rpm and the maximum temperature capability is 200 oe. The mixing bowls are double walled with three independently heated zones. The middle zone is both air cooled and heated to maintain precise temperature control and prevent The temperature of mixing was overheating due to heat dissipation during shear mixing. determined by the viscosity of the system being compounded. The viscosity of the system being compounded determines the shear stresses generated, which in tum determine how effectively the agglomerates are broken and the particle system is dispersed. However, while generating high shear stresses to effectively break up agglomerates, care was taken not to exceed the torque limits of the mixer. Therefore, the temperature of mixing used for a specific system was dependent on the viscosity of the system.

Therefore, the temperature of mixing used for a specific system was dependent on the viscosity of the system. Higher temperatures (100-120°C) were used for systems with high viscosities and lower temperatures (70-90°C) were used for systems with low viscosities [5].The required amount of binder is weighed and added to the mixer and allowed to melt completely. Once molten, dry powder is added in measured increments of 5-40 grams, depending on the viscosity of the binder system. After each addition, the torque of the system increases instantaneously and then slowly decreases and stabilizes over a period of few minutes. The next incremental amount is added after the steady state torque is reached. The process is continued until a desirable powder loading is achieved, after which mixing is continued for a fixed period of time (1-2 hours). The incremental amount of powder additions is decreased at high volume fractions (>45%). The rotor speed is kept at 100 rpm throughout the process.

The incremental amount of powder additions is decreased at high volume fractions (>45%). The rotor speed is kept at 100 rpm throughout the process. Figure 6 shows the spikes in torque values for each incremental addition of silicon nitride powder in the RU 1 binder and subsequent torque stabilization and the final stabilized torque for one hour of mixing. The quality of mix from batch-to-batch of a particular formulation was constant as indicated by reproducible viscosity measurements made on different batches after compounding [5]. The compounded powder/binder system is removed from the mixer, cooled to room temperature, granulated and sieved to >Imm size for feedstock for subsequent viscosity measurement and filament fabrication.454V. Filament FabricationAfter compounding, the next step in preparing the feed material for FDC or FDMet process is to fabricate filaments of nominal diameter 0.070".

Filament FabricationAfter compounding, the next step in preparing the feed material for FDC or FDMet process is to fabricate filaments of nominal diameter 0.070". The control software for the FDMTM system sets the roller speed, and hence the flow rate, for a given s~t of build conditions (road width, s~i~e thickness and liquefier head speed) based on the assumpnon that the feed ~aments are O.O?O m diameter. If the filament diameter is less than 0.070" then the flow rate IS le~s than desrra~le, Such und~rflow results m poor bondmg creating roads with smaller widths and thickne.ss. between adjacent roads and layers or creates VOIds between adJacen! roads an~ layers.. If .the filament diameter is much greater than 0.070" then either the filament IS Il;ot f~d mto the h9l;lef1er, which typically has an opening of 0.073", or it may cause overflow resulnng m poor def~lnon .of fme features in the part.

Therefore, it is impo~t to fab~cate filaments for FD process1J?-g :"lth very controlled diameter (0.070" ± 0.001 "). .It IS als? lffiportant that the filaments malnt~ a circular cross-sectional geometry to be appropnately gnpped and fed by the rollers. In dealmg with ceramic or metal loaded filaments, it is also important to ensure homogeneous distribution of the binder and particle phases in the filaments with no density gradients and porosity.In this study, filaments for FDC/FDMetprocessing were fabricated by two different techniques - capillary extrusion and single screw extrusion. Small quantities of straight, rigid filaments for FD trials were fabricated using an Instron capillary rheometer as a piston extrusion device. The capillary rheometer barrel and plunger are mounted on an Instron mechanical test frame. The barrel was heated to 60° C to 120°C, depending on the material under consideration, and filled with compounded and granulated green ceramic or metal feedstock.

The barrel was heated to 60° C to 120°C, depending on the material under consideration, and filled with compounded and granulated green ceramic or metal feedstock. Extrusion was done by moving the plunger through the barrel at a constant speed of 1mm/minute. A capillary die of 0.070" was used to produce 6"-12" length segments of the filaments with a diameter of 0.070" ± 0.002". The filaments used in this study for FDC and FDMet processing of Si02, PZT, WC-Co, and stainless steel were fabricated using the capillary rheometer approach.Optimized feedstock formulations, such as 55 volume% silicon nitride (pre-treated with 3% oleyl alcohol) in RU9 binder formulation, were used to fabricate continuous and flexible filaments using a continuous single screw extrusion process. A Haake single screw extruder with a 1" diameter screw and LID ratio of20:1 was used [8]. The Haake torque rheometer drive unit, used for compounding feedstock, was used for driving the extruder also [8].

The Haake torque rheometer drive unit, used for compounding feedstock, was used for driving the extruder also [8]. Compounded green feedstock material of particles size in the range of 1.4 mm to 4.0 mm were used as the feed material for extrusion through a die of 0.075" diameter attached to the extruder. Extrusion was done at a screw speed of 10 RPM and the temperature was controlled in four different zones along the length of the screw and barrel. Typically, the temperatures in the feed zone and the die region were kept low at 50°C and 65°C, respectively. The temperatures in the mid-section of the screw length were kept at 75°-100°C. As the filament was extruded, it directly passed onto a moving conveyor. The conveyor speed was adjusted and set to produce controlled drawing of the filaments into final diameter of 0.070" ± 0.002" as it is extruded through a die of 0.075".

The conveyor speed was adjusted and set to produce controlled drawing of the filaments into final diameter of 0.070" ± 0.002" as it is extruded through a die of 0.075". From the conveyor, the fIlament passes through an on-line diameter measuring device, LaserMike, which continuously records the diameter of the filament. The continuously extruding filament, after passing through the LaserMike, were wound onto spools using an automatic winder. The spools used for winding the filaments are similar to those used in commercial FDMTM systems, shown earlier in Figure 3.. The diamete~s of fil~ents fabricated by single screw extrusion as well as capillary extruSIon were conSIstently m the range of 0.070" ± 0.002". Effort is being made to further enhance the diameter uniformity to 0.070" ± 0.0005" by employing pressure control devices, such as a metering pump, in the single screw extrusion process line.

Effort is being made to further enhance the diameter uniformity to 0.070" ± 0.0005" by employing pressure control devices, such as a metering pump, in the single screw extrusion process line. The density measurements made on randomly chosen segments of filaments using He pycnometry consistently indicated 100% theoretical density of the mixed feedstock, suggesting no porosity or density gradients along the455lengths of the filaments. As shown in Figure 7, the microstructure of the cross-section of the green filaments exhibited uniform and homogeneous mixing with no phase separation.VI. ConclusionsThis study describes the various physical, thermal, and mechanical properties required in the filament feedstock for successful FDC and FDMet processing of ceramics and metals using commercial FDMTM systems.

A unique binder series, designated as RU binders, for FDC and FDMet processing has been developed and successfully applied to a wide range of ceramic and It has been shown that through optimization of the RU binder system and metal systems. appropriate selection of dispersant for a specific particulate system, automated FDC and FDMet processing of many particulate ceramic, metal, and composite system is feasible. Appropriate green feedstock compounding and filament fabrication procedures for FD processing have been established. Green feed filaments with controlled diameter and material homogeneity have been successfully fabricated and used for FDC/FDMet processing.Acknowledgmen tsThe work on various material systems have been done through support from DARPA, ONR (Contract No. N00014-94-0115), and the New Jersey Commission on Science and Technology. The authors would like to thank Drs. P.L Whalen and C.P. Ballard of AlliedSignal Inc., Drs. E. Krug and G.

N00014-94-0115), and the New Jersey Commission on Science and Technology. The authors would like to thank Drs. P.L Whalen and C.P. Ballard of AlliedSignal Inc., Drs. E. Krug and G. Carrasquillo of Certech Inc.References1. Proceedings of the Solid Freeform Fabrication Symposium, Vols. 1-7, 1990 - 1996, Edited by H.L. Marcus, et. aI., The University of Texas at Austin, Austin, Texas.2. M.K. Agarwala, et. aI., "Fused Deposition of Ceramics and Metals: An Overview," ibid., Reference #1, Vol. 7, 1996.3. M.K. Agarwala, et. aI., "Structural Ceramics by Fused Deposition of Ceramics," ibid., Reference #1, Vol. 6, 1995, pp.1-8.4. M.K. Agarwala, et. aI., "Fused Deposition of Ceramics for Structural Silicon Nitride Components," ibid., Reference #1, Vol. 7, 1996.5. "Solid Freeform Fabrication of Advanced Ceramics," First Year Annual Report, Oct. 1994 Sept. 1995, Office of Naval Research, Contract No. N00014-94-C-0115.6. U.S. Patent # 5,121,329, June, 1992. 7 . J.W. Comb and W.R.

1994 Sept. 1995, Office of Naval Research, Contract No. N00014-94-C-0115.6. U.S. Patent # 5,121,329, June, 1992. 7 . J.W. Comb and W.R. Priedeman, "Control Parameters and Material Selection Criteria for Rapid Prototyping Systems," ibid., Reference #1, Vol. 1993, pp. 86-9l.8. D.l Shanefield, "Organic Additives and Ceramic Processing," Kulwer Academic Publishers, 1st Edition, 1995, pp. 251-254.9. Haake Fisons Instruments, Paramus, New Jersey.TABLE I Binder Components in the RU Series of BindersBinder ComponentWeight % Component's Role Thermal Degradation Temperature Ran2eRan2ePolymer10 -45Acts as a backbone100 - 510°CElastomer30 - 65Imparts flexibility275 - 500 °CWax15 - 50Viscosity modulator200 - 500°CTackifier10 - 40Promotes adhesion190 - 475°C456.j:::. U1 .......Table II Binder Systems for FDC and FDMetParticulate Binder Vol.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFilled Thermoconductive Plastics for Fused Filament FabricationT. Mulholland1, A. Falke2, N. Rudolph1 1. Department of Mechanical Engineering, University of Wisconsin, Madison, WI 53706 2. RWTH Aachen University, North Rhine-Westphalia, GermanyAbstractAlthough thermoplastics have traditionally been ignored as heat transfer materials, filled thermally conductive plastics fabricated in complex geometries can serve as low-cost, geometrically complicated heat exchangers, along with other thermal applications such as heat sinks.

Thermally conductive, 3D printed heat exchangers can have a significant impact on power generation technology at many scales, with advantages in allowing distributed, on-site manufacturing on multiple, independent machines, and the virtual elimination of scrap. This can command significant cost savings compared to traditional, machined metal heat exchangers. This work examines compounds of polyamide 6 with copper fiber and sphere fillers produced in a co- rotating twin screw extruder. The thermal conductivity is measured and related to the filler content, shape, and orientation.Keywords: 3D Printing, FFF, Thermal ConductivityIntroductionDespite a wide variety of manufacturing processes and a growing list of suitable materials, additive manufacturing (AM) has traditionally been regarded as a rapid prototyping technique.

Several obstacles remain for producing competitive, functional parts.Two major barriers are the anisotropy and weld line formation introduced by the layer- upon-layer processes and their effect on the mechanical properties. This can be seen in studies of build orientation [1-8] and bead orientation [1-3, 5, 6, 9-13]. The part strength is higher in cases where the load is applied in the in-plane direction since it depends on the material itself, and the strength is lower when the load is applied in the build direction since the strength reflects the bonding between layers. Notably, recent studies suggest that the orientation of the individual beads in a part produced by Fused Filament Fabrication may permit strengths almost as high as an injection molded part, if the printed part is sufficiently dense [14].

Fused filament fabrication (FFF), otherwise known by the trademarked name fused deposition modeling (FDMTM), is a widely known and easily accessible AM technique with a huge potential for functionality growth, given the essentially limitless choices of possible materials. However, commercial functional materials are only now starting to enter the market.AM already excels at creating parts with traditionally-impossible, complex geometries, and the potential uses of the technology will continue to grow along with the material choices. The effects of orientation are important for other material properties besides the mechanical strength. Oriented conductivity in AM parts could be a boon for the AM sector in both electrical and thermal applications.871There are many applications that could benefit from oriented conductivity, from electrical conductors and circuits to thermal applications and heat fins.

Importantly, in most of these cases the mechanical strength is less important than other properties. Thus, these applications can serve as a platform from which to grow industrial use of AM technologies, where AM parts can directly compete with traditional manufactured parts.Since polymers have a very low conductivity, filler particles must be added to achieve enhanced thermal or electrical conductivity. The orientation of filler materials in plastic resins during injection molding is known to have a strong influence on the thermal conductivity, and to be a direct result of the manufacturing method [15, 16]. The most important factors influencing thermal conductivity in filled plastics are the filler content, size, and shape. Thermal conductivity is best enhanced when continuous pathways are chained through the part via contact between the filler particles [17].

Thermal conductivity is best enhanced when continuous pathways are chained through the part via contact between the filler particles [17]. The combination of continuous pathways and oriented conductivity suggest that suitable particles should have a high aspect ratio, for example fibers, and that the filler must be added at relatively high volume fractions. For filled FFF filament, early work by Gray et al. [18, 19] demonstrated increased mechanical performance by reinforcing polypropylene with liquid crystalline polymer (LCP) fibrils. Zhong et al. [20] prepared a glass fiber filled ABS filament for a commercial FFF printer. Although improved strength was achieved, plasticizer and compatibilizer were added to realize the flexibility and handleability needed for processing. Masood and coworkers developed a metal-polymer filament designed for rapid tooling of injection molding inserts [21, 22].

Masood and coworkers developed a metal-polymer filament designed for rapid tooling of injection molding inserts [21, 22]. Iron powder (200-500 μm) with volume fractions of up to 40% was added to a polyamide matrix to make FLM filament feedstock capable of processing through a commercial FFF system. In 2012 the first electrical conductive FFF filament was produced with 15 % wt carbon black to print several flexible sensors and capacitive buttons [23]. The same year, Oak Ridge National Laboratory (ORNL) demonstrated printing with short fiber feedstock in public [24] and evaluated the mechanical properties of printed parts [25]. Protoplant Inc., Vancouver, WA, (www.protoplant.com) and the Kickstarter company Functionalize (www.functionalize.com) recently introduced a carbon black or graphite filled filament for the desktop-printer market for increased thermal and electrical conductivity.

To make FFF filament with increased thermal conductivity a viable printing option, an understanding of the complex interaction between matrix and fillers needs to be established, which is dependent on the material properties, the filler size, shape and orientation as well as the manufacturing processes (compounding, extrusion, printing) involved.This work aims to provide a foundation for developing materials useful for fabricating air-to-water heat exchangers produced via FFF. Complicated geometries can be produced via FFF at little added cost, which allows in this case the manufacture of pin fin heat exchangers with complex pin shapes, such as tapered ellipses and air foils. This easily permits the plastic- based heat exchangers to surpass traditional metal heat exchangers with only modest increases in the base material thermal conductivity.

This easily permits the plastic- based heat exchangers to surpass traditional metal heat exchangers with only modest increases in the base material thermal conductivity. It is worth noting that, due to the high convective resistance associated with the air side of the heat exchanger, the material thermal conductivity only needs to be raised to modest values to make an effective heat exchanger. Additionally, the orientation of shaped fillers along the pin axis can significantly increase the heat exchanger performance [26].872In order to be able to tailor thermal properties for any application, it is important to find an accurate way to measure thermally conductive filament. This can be challenging, since even the sample preparation can alter the filler orientation, leading to changes in the thermal conductivity.

This can be challenging, since even the sample preparation can alter the filler orientation, leading to changes in the thermal conductivity. This article outlines the steps taken to produce filled materials for use in FFF, as well as to analyze the effects of filler type and orientation.MethodsMaterial Polyamide 6 (PA6), sometimes referred to by the original trade name Nylon, was chosen for this study because of its high heat resistance and mechanical properties. Ultramid B33-01 (BASF, Germany) with a melting temperature of 220 °C and a density of 1.13 g/cm3 was used. The moisture content at 50 % relative humidity is 2.6 %. Similar PA6 grades from BASF report thermal conductivity and specific heat capacity values of 0.33 W/m-K and 1500 – 1700 J/kg-K, respectively.The material has an intermediate viscosity, which is important for FFF material production. For filament production, a relatively high viscosity is preferable to aid in molten material handling during line startup.

For filament production, a relatively high viscosity is preferable to aid in molten material handling during line startup. In contrast, FFF machines require a relatively low viscosity, due to the low pressures that can be generated in these devices. Additionally, the use of fillers will increase the viscosity. Thus, an intermediate viscosity PA6 is the preferable choice for mixing with the conductive fillers.Spherical copper particles with a nominal diameter below 45 micrometers were purchased (Chemical Store, USA). The spheres are 99.5% copper, with a small iron content and a density of 8.9 g/cm3. They are produced via air atomization of molten copper. 99.9% of the spheres are smaller than 325 µm, with the other 0.1% falling in the range of 45 to 74 µm.Copper fibers with a nominal diameter of 30 µm and a nominal length of 3 mm were purchased from Deutsches Metallfaserwerk (Germany).

The fibers are a minimum 99.9% copper, with a density of 8.9 g/cm3 and a thermal conductivity of 383 W/m-K. To produce the fibers, copper strands are first made into a fleece, which is then chopped.Material Compounding The copper spheres and copper fibers were mixed separately into the PA6 using a twin screw compounding extruder with 27 mm intermeshing, co-rotating screws and a 36 diameter length (927 mm), manufactured by Leistritz Corp. (USA). Twin screw extruders are known to be superior for the mixing of highly filled systems when compared to traditional single screw extruders. Different elements of the fully-customizable screw set are used to melt the resin, generate pressure, and homogenize the melt temperature and filler distribution.

Different elements of the fully-customizable screw set are used to melt the resin, generate pressure, and homogenize the melt temperature and filler distribution. The screw design, along with qualitative descriptions of the screw elements, is shown below in Figure 1.While the PA6 pellets are supplied by a first gravimetric feeder to the extruder at the beginning of the screw, the filler is supplied by a second gravimetric feeder after the resin is melted using twin, intermeshing augers, aptly named the side stuffer. The degassing zone, near the end of the extruder, allows volatiles to escape the material before extrusion, including for873example volatile plasticizers, gaseous degradation products, or water vapor. The latter option is relevant for PA6 processing, since this material is very hygroscopic.Figure 1. Screw design for compounding PA6 with copper spheres or fibers.The parameters listed below in Table 1 were used to produce the copper filled PA6 materials.

Screw design for compounding PA6 with copper spheres or fibers.The parameters listed below in Table 1 were used to produce the copper filled PA6 materials. The copper spheres were produced with a flat temperature zone profile at 240 °C, which resulted in a melt temperature of 253 °C. The copper fiber filled PA6 was produced with a steady temperature zone ramp from 220 to 240 °C, with a melt temperature around 240 °C. The extruder screw speed was 100 RPM, and the side stuffer speed was 140 RPM.Table 1. Twin screw extruder parameters for copper fiber filled PA6.Copper Fiber Volume FractionMelt TemperatureResin Feed RateFiller Feed RateTotal Mass Flow RatePuller Speed20%253 °C6.7 kg/hr13.3 kg/hr20 kg/hr11 m/minThe material was extruded through a three-strand die with an orifice diameter of 3.0 mm. The strands were pulled at a constant velocity by a precision belt puller (ConAir, USA).

The strands were pulled at a constant velocity by a precision belt puller (ConAir, USA). Uneven flow through the die, caused by uneven pressure distribution and gradual clogging of the orifices, caused differences in the material flow rate, which yielded strands of differing diameters. The strands were later classified as D1 (diameter range 1.1 to 1.76 mm), D2 (diameter range 1.8 to 2.4 mm), and D3 (2.44 to 3.3 mm). The different flow rates and the take-up by the belt puller led to natural differences in the fiber orientation, which were subsequently analyzed by different techniques.Thermogravimetric Analysis (TGA)The actual filler content was determined by thermogravimetric analysis (TGA) with a Netzsch TG 209 F1 Libra. In this technique, the sample, about 3 mg of polymer if unfilled, is heated at a constant rate in an inert gas atmosphere, in this case nitrogen, and the weight is constantly measured with a resolution of 0.1 µg.

These samples were heated up to 600 °C at a rate of 20 K/min with a constant nitrogen flow of 20 mL/min. To exclude the weight of water in hygroscopic PA6, the initial weight of the compound was taken as the sample weight at 270 °C. The mass of the filler is taken as the sample weight at 600 °C. These two masses are used to directly calculate the mass fraction of filler, and the densities of each material can then be used to calculate the volume fraction of filler.874Laser Flash Analysis (LFA)LFA is used to determine the thermal diffusivity of a sample. The instrument, a Nanoflash 447 from Netzsch Instruments (Germany), fires a xenon flash lamp at the bottom side of the sample, while a liquid-nitrogen cooled infrared detector monitors the top side of the sample.

The diffusivity of the sample can then be calculated from only the sample thickness (L ) and the half time (t1/2) of the temperature rise on the top surface, according to the following [27]:𝛼 =1.38 𝜋2𝐿2 𝑡1/2Additional models are available to correct for certain errors. This work used the Cowan model [28], a standard choice that corrects for radiative or convective losses from the top and bottom surfaces of the sample, as well as a correction for the time of the energy pulse applied. The samples in this study were measured over a temperature range relevant to the eventual air- to-water heat exchanger, from 25 °C to 105 °C, at 10 K increments.

The samples in this study were measured over a temperature range relevant to the eventual air- to-water heat exchanger, from 25 °C to 105 °C, at 10 K increments. The thermal conductivity (k) can later be directly calculated from the thermal diffusivity (α), the density (ρ), and the heat capacity (cp) by the equation𝑘 = 𝛼𝜌𝑐𝑝The heat capacity is measured over the relevant temperature range using a differentialscanning calorimeter (DSC), DSC 214 Polyma from Netzsch Instruments (Germany).LFA Sample PreparationThe minimum sample size for this instrument is 6 mm round, which makes the measurement 3 mm or smaller filament impossible without some additional sample preparation. Therefore, square samples with 10 mm side length with a recommended thickness around 1 mm were prepared by the techniques described below. Samples were sanded with progressively finer grains to obtain flat, level surfaces.

Samples were sanded with progressively finer grains to obtain flat, level surfaces. They were then spray-coated with graphite, which aids in the absorption of the energy pulse.In order to study the effects of filler orientation on the thermal conductivity, samples were prepared in several different ways, which are depicted below in Figure 2. The filament produced in the compounding process had three different diameters D1 (about 1.2 to 1.6 mm), D2 (about 1.7 to 2.4 mm), and D3 (about 2.5 to 3.3 mm). These were cut into 10 mm lengths, then placed lengthwise into a 10 mm by 10 mm cavity and compression molded to a height of 1 mm, 2 mm, or 3 mm. Smaller filament strands were stacked in order to fill the mold to an adequate height.

Smaller filament strands were stacked in order to fill the mold to an adequate height. Samples measured through the thickness this way demonstrate the through- plane (TP) conductivity, since the conductivity measured is perpendicular to the filament extrusion direction, which is also the dominant direction of fiber orientation.Other samples were fabricated from the first TP samples to measure the in-plane conductivity. The 10 mm x 10 mm TP samples were cut into 1 mm strips, and each strip was875( Eq. 1 )( Eq. 2 )then individually rotated 90 degrees. The rotated strips were placed together in a laminate sample holder (LAM). See Figure 2 for an illustration.Finally, another set of samples was fabricated to measure the in-plane conductivity in a different fashion. The filament produced in the compounding process, with diameters D1, D2, and D3, is cut into lengths slightly longer than 1 mm, 2 mm, or 3 mm.

The filament produced in the compounding process, with diameters D1, D2, and D3, is cut into lengths slightly longer than 1 mm, 2 mm, or 3 mm. These are then placed standing up in the 10 mm x 10 mm mold and compression molded. The length of the filament sections was adjusted iteratively to minimize the material flow while still creating a solid sample. These in-plane compression molded (IPCM) samples also measure the in-plane conductivity. As with the other sample types, the fiber orientation may be affected by the compression.CompressionTPLAMCompressionIPCMFigure 2. Sample production showing filament orientation, which is the principal fiber orientation direction. Top: through-plane (TP) samples are oriented, compressed, then scanned. Middle: in-plane laminate (LAM) samples made from TP samples by cutting and rotating strips.

Top: through-plane (TP) samples are oriented, compressed, then scanned. Middle: in-plane laminate (LAM) samples made from TP samples by cutting and rotating strips. Bottom: in-plane compression molded (IPCM) samples are oriented, compressed, then scanned.Micro-Computed Tomography (µCT) Samples were scanned with a Zeiss (Germany) Metrotom 800 µCT scanner to analyze the fiber orientation. µCT uses an X-ray source to take images of a sample at many different angles, and the data is subsequently reconstructed mathematically into a 3D representation of the sample. Both the filament strands produced on the twin screw extruder and the compression molded samples were analyzed. The scan resolution was 8 µm for the filament strands, and was approximately 20 µm for the compressed samples. This work only judges the fiber orientation qualitatively.Results and DiscussionTGA analysis showed that all samples fell within 1 % vol of the target filler content of 20 % vol.

This work only judges the fiber orientation qualitatively.Results and DiscussionTGA analysis showed that all samples fell within 1 % vol of the target filler content of 20 % vol. The true filler content of each sample was then used to calculate the density of each sample for use in Eq. 2.The specific heat capacity (cp) was measured with DSC on the 20 % vol copper fiber material, and was found to increase linearly (R2 = 0.997) from 0.95 to 1.4 kJ/kg-K over the temperature range from 25 °C to 105 °C.876Before fabricating the samples for the thermal diffusivity measurement, µCT images of the filament were examined and are shown below in Figure 3. The spherical fillers are represented as bright points in the image. The fibers show up as fibers as well as points, which represent out of plane fibers. As expected, the fiber orientation in the extrusion direction is higher in filament of lower diameter.

The fibers show up as fibers as well as points, which represent out of plane fibers. As expected, the fiber orientation in the extrusion direction is higher in filament of lower diameter. The fibers become more oriented in smaller filament strands due to the higher deformation in the flows; smaller cross sections cause higher shear rates which tend to orient the fibers. The fibers are less oriented in the larger diameter filament. This can be seen by the larger number of bright points, as opposed to the fiber shapes.Figure 3. µCT images of the filament strands of sizes D1, D2, and D3 produced by compounding, showing different degrees of orientation at the midplane. From left to right: D2 20% spherical filler, D1 20% fiber, D2 20% fiber, D3 20% fiber.Thermal diffusivity was measured on several different sample types to quantify the effects of compression on the conductivity in the through-plane and in-plane directions. Sample nomenclature is as follows.

Sample nomenclature is as follows. D3 refers to the diameter of the filament used to create the sample. TP, IPCM, and LAM refer to the sample orientation, as explained above. 1 mm, 2 mm, or 3 mm refers to the final thickness of the sample after compression. As an example, D3-IPCM-2mm refers to a sample created with filament of approximately 3 mm in diameter, compressed in the axial direction for measuring in-plane conductivity, and the resulting compressed sample had a thickness of 2 mm.Figure 4 shows the through-plane conductivity of samples made from 2 mm filament compressed perpendicular to the measurement and axial direction. The sample with the least compression, D2-TP-3mm, has the highest through-plane conductivity, between 1.59 and 1.63 W/m-K, which corresponds to the higher number of out of plane fibers visible in the µCT images in Figure 3. Overall the compression ratio has no significant effect.

Overall the compression ratio has no significant effect. The sample with 20 % vol copper spheres shows a significantly lower conductivity, from 0.92 to 1.01 W/m-K.Figure 5 shows the in-plane conductivity for laminate samples (LAM), produced by cutting and rotating the samples from Figure 4, and for in-plane compressed samples (IPCM) produced by axially compressing filament in a mold. It is immediately obvious that the sample preparation methods have a large effect on the conductivity. The best performing laminate sample, D2-LAM-2mm, has a conductivity ranging from 7.04 to 7.74 W/m-K over the temperature range, while the best performing IPCM sample, D2-IPCM-3mm, ranges from 8.48 to 8.95 W/m-K.

Since the fabrication method for LAM samples should tend to raise the in-plane thermal conductivity, and the method for IPCM samples should lower the in-plane conductivity, it can reasonably be expected that the true conductivity of the filament as-manufactured is somewhere between.877y t i v i t c u d n o Cl a m r e h T]K m W/[1.8 1.6 1.4 1.2 1.0 0.8 0.6 0.4 0.2 0.0D2-TP-3mm D2-TP-1mmD2-TP-2mm D2-TP-2mm-Spheres0204060 Temperature [°C]80100Figure 4. Through-plane conductivity in 20%vol fiber-filled samples compressed by different amounts. One sample contains 20 % vol copper spheres.y t i v i t c u d n o Cl a m r e h T]K m W/[8765432D2-LAM-3mm D2-IPCM-3mmD2-LAM-2mm D2-IPCM-2mmD2-LAM-1mm D2-IPCM-1mm100204060 Temperature [°C]80100Figure 5.

In-plane conductivity in 20%vol fiber-filled D2 samples fabricated differently and compressed by different amounts.12y t i v i t c u d n o Cl a m r e h T]K m W/[108642D3-LAM-3mm D3-IPCM-3mmD1-LAM-1mm D1-IPCM-1mm00204060 Temperature [°C]80100Figure 6. In-plane conductivity for LAM and IPCM samples fabricated using large (D3) or small (D1) filament with minimal compression.878120120120Figure 6 shows the effect of different initial filament diameters on in-plane conductivity. D3-LAM-3mm is compared to D1-LAM-1mm, since both samples represent the minimum compression for those filament sizes. Similarly, D3-IPCM-3mm is compared to D1-IPCM-1mm, since both have low compression. There is very little difference in the conductivity of the IPCM samples, while the LAM samples show a difference of about 2 W/m-K.

There is very little difference in the conductivity of the IPCM samples, while the LAM samples show a difference of about 2 W/m-K. However, the conductivity of LAM samples is approximately twice as large as that of the IPCM samples.The effects of changing fiber orientation can be clearly seen in the thermal conductivity results. The differences can also be seen in µCT scans of the compression molded samples. Figure 7 below shows the fiber orientation in three samples made from filament of approximately 3 mm (D3) and compressed by different amounts. Since the fibers initially occupied the mold completely in the filament’s axial direction, the flow during compression is mostly perpendicular to the filament axis, as the material fills in the gaps between the cylindrical surfaces. The fourth image shows a sample comprised of filament compressed axially for measurement of the in-plane properties (IPCM).

The fourth image shows a sample comprised of filament compressed axially for measurement of the in-plane properties (IPCM). In this case, the flow is again perpendicular to the filament axis, which is in the plane of the image in Figure 7. Since the majority of the fibers are still oriented with the filament axis the image below shows most of the fibers as points.It can be generally observed that greater compression leads to a bigger change in thefilament orientation. This corresponds to the thermal conductivity results discussed above.ConclusionSeveral characterization methods were explored for measuring the thermal conductivity of extruded PA6 filament with 20 % vol copper fiber or sphere fillers. It is clear from the thermal measurements, and confirmed with the µCT scans, that the conductivity of the fiber filled samples is strongly dependent on the fiber orientation, as expected. It is shown that the sample fabrication method also strongly affects the conductivity.

It is shown that the sample fabrication method also strongly affects the conductivity. However, the methods were successful in characterizing the range of values for the thermal conductivity in the original filament. It is difficult to give an accurate value for the in-plane conductivity in the unaltered filament, but it should lie in the range between the in-plane compressed samples and the laminate samples, from 4 to 10 W/m-K at 25 °C. The through-plane conductivity doesn’t change greatly with compression, so the true value in the filament should lie close to the measured values, approximately 3.5 W/m-K.These values can be used for the design and simulation of the 3D-printable heat exchanger. Since the printing direction in FFF corresponds to the high in-plane conductivity, this material is advantageous for pin-type heat exchanges, since these thin structures are normally printed parallel to the pin axis.

Simulations and testing are needed to determine the overall effectiveness of the heat exchanger, especially considering that the walls of the heat exchanger should correspond to the measured through-plane conductivity values. However, it is worth noting that 20 % vol fibers perform twice as well as 20 % vol spheres in the through-plane direction.879Figure 7. µCT images showing varying degrees of orientation in compressed samples at the midplane through the thickness. Top left: D3 TP 3 mm sample, top right: D3 TP 2mm sample, bottom left: D3 TP 1 mm sample, bottom right: D3 IPCM 1 mm sampleFuture WorkMore tests are currently underway to compare these results with different filler volume contents. This may provide different results, considering that higher volume contents cause more fiber jamming, which leads to lower fiber orientation. Rheometry will be used to study the interaction between filler volume content and shape and the melt viscosity.

Rheometry will be used to study the interaction between filler volume content and shape and the melt viscosity. The µCT data is also being used to generate orientation matrices, which will be helpful in drawing quantitative relationships between the orientation and the thermal conductivity. These orientation matrices are derived from a method developed in-house at the Polymer Engineering Center at the University of Wisconsin – Madison [29].880Work is also ongoing on modifying traditional FFF equipment to be able to handle highly conductive materials. Possible problems may include buckling and breaking near the extruder motor, or softening there due to heat transfer from the nozzle tip along the highly conductive fiber direction. Additionally, high metallic filler content may lead to jamming and clogging of the nozzle, which could require a redesign of the nozzle geometry.Other filler materials should also be investigated.

Additionally, high metallic filler content may lead to jamming and clogging of the nozzle, which could require a redesign of the nozzle geometry.Other filler materials should also be investigated. Carbon fiber, for example, is more conductive and lightweight, which is important for mobile heat transfer applications. The materials developed through this work will be tailored for industries such as refrigeration, power electronics, and automotive.AcknowledgementsThe Polymer Engineering Center at the University of Wisconsin – Madison and the authors would like to thank BASF for material, Leistritz for the donation of the twin-screw extruder, and the U.S. Department of Energy ARPA-E for funding.References[1] J. C. Riddick, M. A. Haile, R. Von Wahlde, D. P. Cole, O. Bamiduro, and T. E. Johnson, “Fractographic analysis of tensile failure of acrylonitrile-butadiene-styrene fabricated by fused deposition modeling,” Addit. Manuf., vol. 11, pp. 49–59, 2016.[2] A. M.

Film Fabrication Process on Transparent Substrate using Mask Projection Micro- StereolithographyAmit S. Jariwala, Fei Ding, Xiayun Zhao, David W. Rosen*George W. Woodruff School of Mechanical Engineering Georgia Institute of Technology Atlanta, Georgia, 30332 *Corresponding author. Tel.: +1 404 894 9668 Email: david.rosen@me.gatech.edu Reviewed, accepted September 10, 2008 Abstract:In this study, a Mask Projection Micro-Stereolithography (MPµSLA) process with the ability to cure a film of various thicknesses on transparent substrates is presented. Incident radiation, patterned by a dynamic mask, passes through a transparent substrate to cure photopolymer resin layers that grow progressively from the substrate surface. When compared to existing Stereolithography techniques, this technique eliminates the necessity of recoating, reducing process time and improving accuracy. A film of varying thicknesses can be fabricated on flat or curved transparent substrates.

A film of varying thicknesses can be fabricated on flat or curved transparent substrates. Models of the optical system and resin cure are developed and reported. An existing MPµSLA process planning method is being extended to account for radiation transmission through a substrate. The models are verified using experiments.1. IntroductionMask Projection Stereolithography (MPSLA) is an additive fabrication process used to build physical components out of a photopolymer resin. The CAD model of the part to be built is sliced by horizontal planes and the slices are stored as bitmaps. These bitmaps are displayed on a dynamic mask and are imaged onto the photopolymer resin surface. When a bitmap is imaged onto the resin surface, a layer corresponding to the shape of the bitmap gets cured. This layer is coated with a fresh layer of resin by lowering it inside a vat holding the resin and the next layer is cured on top of it. By curing layers one over the other, the entire part is built.

This layer is coated with a fresh layer of resin by lowering it inside a vat holding the resin and the next layer is cured on top of it. By curing layers one over the other, the entire part is built. This technology has been demonstrated in various papers, like Bertsch et al., (1997), Chatwin, (1998), Monneret et al., (1999), Sun et al.(2005) and Limaye and Rosen, (2007). Commercially available machines include Perfactory® range of machines from EnvisonTec, Germany. The principle behind these machines is similar to that mentioned in the papers above, and sometimes employs more than a single DMD chip to produce multiple parts.

The principle behind these machines is similar to that mentioned in the papers above, and sometimes employs more than a single DMD chip to produce multiple parts. One of the interesting aspects of these machines is that the irradiation from the DMD chip passes through a transparent substrate into the resin vat, compared to irradiance from the top of the vat as in conventional micro-SLA processes.The technologies mentioned above cure parts in a layer by layer fashion on a flat substrate by irradiance from the top of the resin surface. Research has been done on curing parts through transparent substrates, which may enable thin film coating, deposition of micro-channel or modifying a surface of an optical part. Erdmann et al. (2005) had shown the use of mask projection SL through transparent substrates for manufacturing of micro-lens arrays. Mizukami et al.

Erdmann et al. (2005) had shown the use of mask projection SL through transparent substrates for manufacturing of micro-lens arrays. Mizukami et al. (2002) had proposed the use of curing by laser beam scanning through transparent substrates for manufacture of micro-electrophoretic chips.In this paper, a film fabrication process on transparent substrate using Mask Projection Micro Stereolithography is introduced. The laser beam patterned by the dynamic mask is projected on the transparent substrate to cure the resin. A process planning method is proposed to control the bitmaps and the time of exposure projected on the dynamic mask for each bitmap. This helps to achieve accurate control over the surface of the cured film. This process can be used to fabricate films on flat or curved substrates.The experimental setup is discussed in Section 2. Apart from the desired model of the part, the inputs required for process planning include an irradiance model and resin characterization.

Apart from the desired model of the part, the inputs required for process planning include an irradiance model and resin characterization. The irradiance model provides a relationship between the micro-mirrors displayed on the DMD surface and the radiant energy received on the transparent substrate. Resin characterization data relate the response of the photopolymer resin to the irradiated energy by means of a working curve. This working curve helps to determine the energy required to be irradiated on the transparent substrate for a given part geometry. Both the irradiance model and the resin characterization are explained in Section 3.In Section 4, the overall process planning method is explained. The analytical models for resin characterization and irradiance distribution (obtained in section 3) are fed into a regression model to establish the time of exposure for which each micro-mirror should be displayed in the ‘ON’ state.

In order to optimize the process for displaying the micro-mirrors, a bitmap generation algorithm is used to group the micro-mirrors which need the same time of exposure.Sample film components are assumed to be designed by the proposed process plan. The bitmaps and the corresponding time of exposure are calculated for these cases using the proposed approach. These simulation results and experimental verification are presented in Section 5.2. Experimental Setup The schematic the MPµSLA of system developed is illustrated in Figure 1. The specifications of the are system presented in Table 1. the The design of system can be divided into three modules: conditioning Beam module: This module consists of a UV laser light from source (now, Omnichrome Melles Griot) (Model # 3074-M-X04). The laser emits 38.5mWLaserEngineered diffuserCollimating lensResin vatTransparent substrateImaging lensDMDComputerFigure 1 Schematic of the MPSLA systemTEM01 at wavelength of 325nm.

An Engineered™ Diffuser (micro lens array) is placed right after the laser source to homogenize the beam’s intensity profile and enlarge the beam from a diameter of 1.5mm to 50 mm. A UV transmitting Plano-Convex lens with EFL 150.0 mm is used to collimate the light emerging from the diffuser.

The DMD is an array of individually addressable, bi stable micro mirrors, which can be selectively oriented, to display any bitmap. Every pixel on the bitmap controls one and only one micromirror on the DMD. The micromirrors are 12.65 µm square and the spacing between adjacent micromirrors is 1µm. The micromirrors in their neutral state are parallel to the DMD chip. In its “ON” state, a micromirror swivels about its diagonal by 12° in one direction and in the “OFF” state, swivels by the same amount in the opposite direction. The DMDTM is a product of Texas Instruments and was sold by Productivity Systems Incorporated (PSITM). The bitmap displayed on the DMD serves as the object for the imaging system. The bitmap is imaged onto the substrate by the imaging lens. The DMD is mounted parallel to the horizontal plane.

The bitmap displayed on the DMD serves as the object for the imaging system. The bitmap is imaged onto the substrate by the imaging lens. The DMD is mounted parallel to the horizontal plane. Theobject distance, as measured from the center of the pattern to the mid plane of the imaging lens, was 152 mm and the image distance, measured from the mid plane of the imaging lens to the resin surface, was 132 mm. The radius of curvature of the Plano-Convex imaging lens is 34.25 mm and the thickness of the lens is 4.4 mm. The refractive index of the lens material is 1.460 (air). The above-mentioned data was used in the ray-tracing algorithm, described in the section “Irradiance Model”.Resin vat: The resin vat is a rectangular container with the base made of a transparent glass slide. This glass slide acts as the substrate over which the film is cured.Substrate meshed with x points pr1, pr2, …prx, with each point surrounded by an infinitesimal area3.

This glass slide acts as the substrate over which the film is cured.Substrate meshed with x points pr1, pr2, …prx, with each point surrounded by an infinitesimal area3. Analytical modeling this section, is presented produced3. Analytical modeling this section, is presented producedirradiance the model irradiation individual micromirrors on the DMD mask onto the resin surface is computed by first order ray tracing. The working curve for the resin is characterized on this setup using experiments. This working curve is used further as an input for process planning.IntheIMAGING LENSPattern meshed with n points; p1, p2, …pn; where n(cid:1) 8m rays generated from each point in directions v1, v2, …vm; m(cid:1)8DMDIrradiance modelFigure 2 Schematic of ray tracing algorithmIrradiance Model models the irradiance received by the resin in terms of the process parameters.

The irradiance distribution on the resin depends upon the power distribution across the light beam incident on the bitmap and upon the optical aberrations caused by the imaging lens. The ray tracing algorithm is adopted from Limaye & Rosen (2007). Figure 2 shows the schematic of the ray tracing algorithm for projection of light rays from DMD onto flat substrate.The irradiance distribution across the beam incident on the DMD is assumed to be uniform and the value is measured using a radiometer. This irradiance is one of the inputs to the Irradiance model.The Irradiance, H(pri) at a point in the resin can be given by:H(pri) = cnm∑∑d (p v pr i k j ,,)(1)j= 1k= 1where c is a constant. d is a function introduced to evaluate whether a particular ray will strike an infinitesimal area centered on a given point on the resin or not.

d is a function introduced to evaluate whether a particular ray will strike an infinitesimal area centered on a given point on the resin or not. pj corresponds to the number of points: p1, p2,….pn, where n(cid:1) ¥ ) represents the direction vector in which the rays are emitted from the point on the DMD and pri (pr1, pr2,…., prx, where x(cid:1) ¥ ) is a point on the substrate. Since a collimating lens is used in conjunction with the diffuser, the light beam incident on the DMD is fairly collimated with a divergence angle of less than 1 degree. To take into account the effect of the minor divergence, a cone of rays is emitted from each pattern point.on the DMD. vk (v1, v2, …vm, where m(cid:1) ¥Resin Characterization=, where t is the exposure time at this point.

vk (v1, v2, …vm, where m(cid:1) ¥Resin Characterization=, where t is the exposure time at this point. According to a threshold model proposed by Jacobs (1992), a resin point is cured if and only if the exposure received by this point is greater than the threshold exposure of polymerizationFor a point(ipr , exposure)iprE ()is given by( prE i)( prE i)t(cE . )The variation in exposure with depth in the resin follows the Beer Lambert’s law ofabsorption. So, the exposure at a height z in the resin is given as:( prE i), z=( prE i)e/ pDz(2)wherepD is the depth of penetration of the resin. Again, ifprE ( i‡), zEc, the resin will cure atthat point. So, the depth to which the resin will cure at a point when exposed to irradiation for a time t , is given by:ipr receiving irradianceiprH ()Cd(pr i)=Dpln(EtprH i(/)c)(3)The model in Eq.

(3) is based on an assumption that the attenuation of radiation through a cured layer is the same as that through uncured resin. It does not count the effects of radiation through a cured part, which is in solid phase. Limaye & Rosen (2007) have observed experimentally that the attenuation through a cured layer is significantly less than that through pSD is expected to be different the liquid resin. Thus, the depth of penetration for a cured layer pLD . The layer cure model developed by Limaye & Rosen (2007)from that for the liquid resin by modeling the layer curing as a transient phenomenon is given below,dz=DpL1ln[+dEexp(EcDz /pS)](4)After applying Taylor series expansion and omitting the higher order terms, Eq.

(4) can be further simplified asexp(- E The layer curing model is obtained after solving the ordinary differential equation above, D(cid:215)Dz /)dE(cid:215)pS»DdzpLcDE E(cid:215)+pL»(cid:215)pL1)ln(DzpSDDpScpS(5)(6)Through experimentation, the resin cure characteristics were obtained to provide a relationship between the amount of exposure and the curing depth. The resin used with the MPµSLA system under consideration is the Vantico Huntsman SL-5510 resin. The values of Ec and Dp have been specified by the resin manufacturer to be 8.9 mJ/cm2 and 0.122 mm respectively. Research on MPµSLA systems has shown that the experimentally observed values of Ec and Dp differ from their values specified by the manufacturer (Bertsch et al., 2000, Farsari et al., 2000, Hadipoespito, 2003). So, the resin needs to be characterized experimentally to determine the values of Ec and Dp. Fig.

So, the resin needs to be characterized experimentally to determine the values of Ec and Dp. Fig. 3 shows the working curve for the resin obtained by our proposed setup.,following experiments were performed to determine the values of Ec and Dp. A thin film is cured on a flat glass substrate by exposing it to radiation for different time durations. By varying the time of exposure, the radiant energy received by the film is varied. The thickness of the cured film is plotted against the exposure received by the film as shown in Figure 3. From the plot, the value of Ec, DpS and DpL are 4.0mJ/cm2, be found 0.015mm 0.011mm, and respectively, which are lower than manufacturer’s specified value.ThetotheFigure 3 Working curve of Huntsman 5510 resin with the MPµSLA system4.

Process planning methodCAD modelIrradiation modelA process planning method is formulated to generate the bitmaps to be displayed on the DMD and the times for which they should be imaged onto the resin to cure a thick film required with bitmap dimensions. A displayed on the DMD is nothing but a cluster of micromirrors oriented in a particular direction. Using the Irradiance Model, the correspondence between the location of an “ON” micromirror on the DMD and the locations of pixels irradiated by it on the resin can be established. By applying a regression model, the time for which each micromirror needs to be imaged onto the resin to cure a certain part can be computed. The micromirrors with similar times of exposure are grouped together to generate multiple bitmaps.

The micromirrors with similar times of exposure are grouped together to generate multiple bitmaps. Thus, the input to the process planning method is the three dimensionalResin working curveDesired energy profile (Ej)Irradiation from micromirror (Hij)Regression modeltheExposure for each micromirror (Ti)Bitmap generation algorithmTime of exposure for bitmapsBitmaps onDMDFigure 4 Flow-chart of proposed process plangeometric profile of the part to be cured on a transparent flat/curved substrate and the outputs are the bitmaps to be displayed on the DMD and the times of exposure (TOE) for which they should be imaged onto the resin surface. The schematic of the process plan is explained in Fig. 4.The detailed idea of the proposed process planning method could be explained in the context of Figure 5. Fig. 5 shows the CAD model representation of a film to be cured on the substrate.

4.The detailed idea of the proposed process planning method could be explained in the context of Figure 5. Fig. 5 shows the CAD model representation of a film to be cured on the substrate. The base surface area of this CAD model is discretized into a number of square pixels ijZ . Each column with average height, corresponds to one pixel on the substrate. The ijZ along this energy required to cure the height, column can be obtained resin the from ijE . The energy characterization curve as, irradiated on the pixels of the resin is a result of irradiance produced by the micromirrors and the micromirrors are the switched “ON”. For each of the pixels on the resin, the objective is to minimize the deviation of the exposure received from the micromirrors ijE.

For each of the pixels on the resin, the objective is to minimize the deviation of the exposure received from the micromirrors ijE . The time of and the required energy, exposure for each micromirror can be obtained by satisfying the least squares minimization.zCAD modelijZMjtime for which0iFigure 5 Stereolithography parameter estimation formulation exampleNRegression Model: For the sake of this study, irradiance produced by a micromirror on a ijklH , where ij corresponds to the pixels on the substrate (400 x pixel in resin is represented as 400 pixels) and kl corresponds to the number of micromirrors (300 x 300 mirrors) on the DMD ijklH represents the irradiance on the (i x j)th pixel of the flat substrate caused by the mask. Thus, (k x l)th micromirror on the DMD mask. ijklH could be obtained from the irradiance model in Section 3.The parameter optimization algorithm proposed by Sager & Rosen (2008) is adopted to develop the regression model.

ijklH could be obtained from the irradiance model in Section 3.The parameter optimization algorithm proposed by Sager & Rosen (2008) is adopted to develop the regression model. The exposure at each (i x j)th pixel on the resin surface is, klT is the time of exposure of each micromirror.

The exposure at each (i x j)th pixel on the resin surface is, klT is the time of exposure of each micromirror. If the exposure required at each H *ijkl T pixel is, where kl ijE then, the least-squares fitting problem can be formulated as follows,Given: Geometry of the partyxGz =,().Resin properties: penetration depths (for solid and liquid),pD , critical exposure,cE .Irradiance on the (i x j)th pixel on the resin surface by (k x l)th micromirror,ijklH(Obtained from the Irradiance Model).Find: Exposure time for every micromirror:klT , k = 1, 2 …P (number of elements onDMD along x-direction), l = 1, 2 …Q (number of elements on DMD along y-direction).Minimize: minS=NM∑∑i= 1j= 1( E ij~ E ij) ∑∑NM2=i= 1j= 1  EijQP∑∑k= 1l= 1Hijkl~ T kl  Constraints:klT0‡klT , is obtained and klT values is divided into groups, in a micromirror grouping algorithm is applied.

The set of which the values for times of exposure are similar. The average time of exposure for each group, iT , i = 1, 2, 3, … N ( number of groups), is assigned as times of exposure for all the micormirrors in this group. In order to ensure that each pixel on the substrate get continuous curing till desired thickness, the following bitmap generation method is developed.Bitmap Generation: After the time of exposure for each micromirror,1)T 1={ i i TT 1 1=} i iT ,=,...,3,2,1N.t1=min(T 1)Bitmap pattern 1: { = i i TT 2 2= 22)Ti T 1iT 1P =group of micromirrors with 1 min(} .=ttT 2)12t 1Bitmap pattern 2: . . . K) T k2P =group of micromirrors with} 1 , Bitmap pattern K : KP =group of micromirrors with . . . Display each bitmap pattern KP with time duration of Kt{ i i T T k k=== ki T kmin(t)tT kk1iT 2i T kt2tk.In this paper, the bitmap generation process was done manually.

An automatic methodusing kmeans clustering algorithm is under development.5. Results and DiscussionsIn order to experimentally verify the process planning method in the MPµSLA system,three cases are studied.Case 1: The CAD model for Case 1 is shown in Fig. 6(a). The part is composed of two square blocks with a step height of 30 µm. The front-side view of the part is shown in Figure 6(b). To reduce the computation time, a 2D profile in Fig. 6(b), which can be extruded along the y axis to yield the 3D part, is considered in the regression model. From the working curve in Figure 3, the desired exposure profile is plotted in Figure 6(c).The irradiation on the substrate was obtained for every micromirror on the DMD using the irradiance model. Figure 7 shows the irradiance distributions on the substrate corresponding to the micromirror at the DMD center and the micromirror at the middle of an edge. In Figure 7(a), the center area with most intensive irradiation is 10*10 µm.

In Figure 7(a), the center area with most intensive irradiation is 10*10 µm. The size of surrounding area is 20 µm with less than half of the irradiation at the center. For the edge micromirror shown in(7)(8)(9)(10)(11)Figure 7(b), the area with the most intensive irradiation is 10*20 µm. However, because of the offset of the micromirror to the optical axis of the system, more blur area appears on the irradiation area from the edge micromirror than that from the center micromirror.zyx2000 um1000 um30 um30 umAfter the irradiation for each micromirror on DMD was generated, ijklH , the contribution of irradiation on the (i x j) th pixel on the substrate from the (k x l)th micromirror was obtained from the corresponding irradiation profile directly. Then, the regression model in Eq. (7) was solved by linear least squares method to the time of exposure for every individual micromirror. After the regression model was solved, the bitmap generation algorithm (Eq.

(7) was solved by linear least squares method to the time of exposure for every individual micromirror. After the regression model was solved, the bitmap generation algorithm (Eq. 8-10) was applied to produce the sequence of bitmaps and the time of1l=30 um(a) exposure for each bitmap. Figure 8 shows the two bitmaps which were displayed on the DMD for the specified time in order to cure the desired shape as shown in Figure 6(a) for Case 1. Figure 8(a) is the first bitmap displayed on DMD. After the first bitmap was displayed on DMD for 47 seconds, it was replaced==2l3l1000 um1000 um=4l60 um2000 um(b))W m( n o i t a d a r r Ii201510Desired energy profile50100050005001000X (um)(c) Figure 6 (a) The CAD model (desired cured shape) for Case 1; (b) The front-side view of the desired cured shape for Case 1; (c) The corresponding desired irradiation profile for Case 1.by the second bitmap shown in Figure 8(b).

The second bitmap was displayed on the DMD for 111 seconds to finish the curing process.The cured shape from the experiments for Case 1 is shown in Figure 9. The area enclosed by the blue lines is the cured part. Since the film is cured on a transparent substrate, measurements for thickness are performed from the base of the glass substrate. The hazy images on the other side of the measured datum are from the reflections produced from the transparent substrate. Comparison of the cured part and the desired shape in Fig. 6(b) is shown in Table 2. The cured shape from the experiment agrees with the shape from the simulations.

The 2D profile of the CAD model is shown in Figure 10(b). After applying the regression model, the bitmap series54 µmdisplayed on the DMD are listed in Fig. 11. Figure 12 shows the cured part from the experiments.

Figures 14 and 15show the bitmaps series for the DMD and the experimental results, respectively. All the dimensional errors in Case 2 and Case 3 are less than 10%.(a) 1st bitmap (T=39s)(b) 2nd bitmap (T=36s)(c) 3rd bitmap (T=66s)(d) 4th bitmap (T=103s)(e) 5th bitmap (T=149s)(f) 6th bitmap (T=189s)(g) 7th bitmap (T=208s)(h) 8th bitmap (T=215s)(i) 9th bitmap (T=181s)(j) 10th bitmap (T=213s)Figure 14 Bitmaps displayed on DMD for Case 362 µmRadius: 7530 µm2004 µmFigure 15 Cured shape from experiments for Case 3.5. ConclusionsIn this paper, Mask Projection Stereolithography was demonstrated as an applicable process for fabrication of films of varying thickness on transparent substrates. A detailed process plan based on the Energy Threshold model was presented.

A detailed process plan based on the Energy Threshold model was presented. The proposed process planning method provided the bitmap sequence and time of exposure for each bitmap to be displayed on the DMD mask.The results from the process planning method enabled curing of a film with the desired profile. However, the experimental results suggest the necessity to further improve the process planning method by incorporating additional constraints in the regression model like, resin cure kinetics, oxygen inhibition effect, shrinkage, etc. These factors would be incorporated in our future work to gain a better control over the generated cure profile of the film. Extension of the proposed fabrication process for different resins shall be studied for increasing the potential applications of this process.6. ReferencesBertsch A., Zissi S., Jezequel J., Corbel S., Andre J.

nishing of ABS-M30 Parts Manufactured with Fused Deposition Modeling With Focus on Dimensional AccuracyM. Fischer1,2; V. Schöppner11Kunststofftechnik Paderborn (KTP), University of Paderborn, D-33102 Paderborn, Germany 2Direct Manufacturing Research Center (DMRC), University of Paderborn, D-33102 Paderborn, GermanyAbstractFused Deposition Modeling (FDM) parts are prone to process-related rough and wavy surfaces with stair-stepping effects whenever the parts produced have sloped or rounded geometries. These stair-stepping effects can be reduced by using a smaller slice height, but complete elimination is not possible. In this paper, FDM parts manufactured with the material ABS-M30 are finished using mass finishing methods. The mass finishing is done with a trough vibrator, which is comparatively gentle to the parts in comparison to other mass finishing technologies.

The mass finishing is done with a trough vibrator, which is comparatively gentle to the parts in comparison to other mass finishing technologies. The analysis discusses the surface-smoothing effect of finishing time and intensity on various part sizes and build orientations. In addition, the dimensional accuracy of the parts after the finishing process is examined.IntroductionAdditive manufacturing processes make it possible to directly convert a 3D digital CAD model into a physical component. In contrast to formative or subtractive manufacturing processes, which create the desired shape by re-forming the initial material or by removing material from the initial form, additive manufacturing produces components by repeatedly adding volume elements to the part being built. This technology enables rapid production of prototypes and offers designers new creative freedom when developing end products.

This technology enables rapid production of prototypes and offers designers new creative freedom when developing end products. Fused Deposition Modeling (FDM), which belongs to the category of extrusion processes, offers one possibility of employing additive manufacturing. [1] [2]The process steps necessary when using FDM units such as those from the company Stratasys can be divided into three areas: data preparation (creating a build file), job preparation (prepping the machine), and the building process itself on the FDM unit. The 3D CAD model, available as a .stl file, is sliced into layers using the build-preparation software Insight. This software also calculates any necessary supporting structures and the tool path parameters. Afterwards, the prepared data is sent by the software to the control center, which then determines the position of the part within the build chamber.

Afterwards, the prepared data is sent by the software to the control center, which then determines the position of the part within the build chamber. The control center then sends the preprocessed and calculated data, packaged as one “print” job, to the FDM machine unit.The initial building material is a spool of thermoplastic polymer thread which has been placed in the machine. This support material is distinct from the material which is used later to build the part; the extrusion head switches between the two materials during the build process. A schematic representation of the FDM process is shown in Figure 1.zyxModel MaterialSupport MaterialBuildSheetFDM HeadFigure 1:FDM ProcessBoth the part material and the support material are fed into the heated FDM extrusion head. In the extrusion head, the material is liquefied and pushed out through an extrusion tip.

In the extrusion head, the material is liquefied and pushed out through an extrusion tip. The material is applied at defined locations in 2-dimensional space by the extrusion tip and by the FDM head, which can move in the X and Y directions. After completing one part layer by applying multiple adjacent ribbons of melted material, the build platform is lowered in Z- direction to make space for the next layer; this process is repeated until all layers of the part have been built. Any support material can be removed after the build process, either by simply being broken off or by being dissolved in a lye solution. [2]In addition to its many advantages, such as the production of complex part structures, and the ability to dispense with product-specific tools for small-series production, Fused Deposition Modelling nevertheless offers many process-related challenges as well. These include the geometric exactness of the parts produced, and in particular the quality of part surfaces.

These include the geometric exactness of the parts produced, and in particular the quality of part surfaces. As a result of the layer-by-layer build process and the placing of adjacent polymer strands, parts produced using FDM exhibit rough and wavy surface structures with stair-stepping effects at edges and corners [4]. The FDM machine manufacturer Stratasys offers the possibility of producing parts from, among other materials, ABS-M30 with varying layer heights of 0.127 mm (0.005 in.), 0.178 mm (0.007 in.), 0.254 mm (0.010 in.), and 0.330 mm (0.013 in.) when using the Fortus 360mc, 400mc, or 900mc printer units. Each layer thickness requires a corresponding tip diameter, which must be consistent throughout the print job due to system constraints. The tip diameters decrease as the layer thickness decreases and have labels ranging from T20 (for up to 0.330 mm layer thickness), to T16, T12, and lastly down to T10 (for 0.127 mm layer thickness).

The tip diameters decrease as the layer thickness decreases and have labels ranging from T20 (for up to 0.330 mm layer thickness), to T16, T12, and lastly down to T10 (for 0.127 mm layer thickness). By increasing layer thickness, the build time can be shortened; however, this also leads to more prominent stair-stepping effects, which in turn reduces the exactness of the surface structure. [3] [13]In order to improve the visual aspects of parts, post-processing of the surface should be carried out on the part. Stratasys offers a chemical post-processing option for this purpose by the use of the Finishing Touch Smoothing Station, with which the surface of parts from the ABS family can be smoothed.

Stratasys offers a chemical post-processing option for this purpose by the use of the Finishing Touch Smoothing Station, with which the surface of parts from the ABS family can be smoothed. Should chemical post-processing of the parts be undesirable, which is quite possible depending on the eventual application, mechanical finishing processes likevibratory grinding or abrasive blasting offer possibilities of improving part surface quality; this paper will examine the influence of vibratory finishing processes on the surface quality of ABS- M30 parts produced by FDM. In addition, there is a short discussion of abrasive blasting; however, this process is not discussed in detail here due to problems with reproducibility.Materials and MethodologyFDM parts with layer thicknesses of 0.127 mm (T10 tip) and 0.178 mm (T12 tip) were produced for this purpose using a Fortus 400 mc 3D FDM printing unit from Stratasys.

The following vibratory grinding process was carried out using an RMO 210/530 TE-30 trough vibrator from the company RÖSLER Oberflächentechnik GmbH. The abrasive blasting was performed with a blasting station from the manufacturer Normfinish, type JUNIOR EP 0275, with glass spheres (diameter 70-110µm) as the finishing media. Furthermore, the blasting tests were carried out by hand, meaning that results were strongly dependent up on the available work force; thus they should only be used for a general comparison. For the examination of the surface structure as well as the dimensional accuracy of the parts, a mechanical profilometer from Jenoptik (model: Hommel ETAMIC T8000) was used in addition to a Colibri Multi 3D optical measuring system from Fraunhofer Vision. [5] [10]Surface Quality of FDM-Produced Parts of ABS-M30Due to the production process, parts produced in a FDM process display notably anisotropic mechanical properties [6].

[5] [10]Surface Quality of FDM-Produced Parts of ABS-M30Due to the production process, parts produced in a FDM process display notably anisotropic mechanical properties [6]. Analogously to this quality of the mechanical properties, the surface characteristics also exhibit differences with respect to the build and the measurement directions [8]. Differing surface qualities are obtained for different sections of the part, which depend on the build process parameters, contact to the supporting material, and the orientation of the part itself. Characterization of the surface quality can be done by various measuring processes, such as different optical or mechanical methods. One mechanical method is the use of a stylus-equipped contact profilometer; important characteristic values, such as the average surface roughness Rz, can hereby be determined for a two-dimensional surface profile according to DIN EN ISO 4287.

[7] [14]In the following, the test specimens of ABS-M30 are produced with the respective Standard Tool Path parameters according to the Insight 9.0 software and with the function “Visible Surface Style: Enhanced”. Further, specimens with a layer thickness of 0.127 mm (T10 tip) and those with a thickness of 0.178 mm (T12 tip) are considered separately for purposes of comparison.

Further, specimens with a layer thickness of 0.127 mm (T10 tip) and those with a thickness of 0.178 mm (T12 tip) are considered separately for purposes of comparison. An overview of the measured average surface roughness values Rz is shown in Figure 2 for varying layer thickness and build orientation.Rz/ µmT12020406080100120140T10(Support Contact)BuildangleMeasuringperpendicularTop SideMeasuringparallelUndersideX (0 )YZ (90 )SpecimenFigure 2:Average Height of Profile Rz for Different Slice Heights (0.127mm/T10 tip and 0.178mm/T12 tip), Build Orientations and Part SurfacesThe schematic representation of the specimen in Figure 2 shows a complex relationship between the measuring direction, the side of the part which is measured, and the build orientation of the surface being measured.

The layer-by-layer building technique, typical of FDM, is regarded here as a periodic profile, for which, according to DIN EN ISO 4288 surface measurements should be carried out perpendicular to the furrows created during production in order to obtain reliable measuring results. This entails measuring perpendicular to the layer structure, which is shown in Figure 2 as “Measuring Perpendicular”. In the bar chart in Figure 2, significant differences in surface quality with relation to build orientation and layer thickness are clearly visible for specimens of dimensions of 75x50x20 mm, which have not been post- processed. For a build orientation of 45° in particular, which is usually heavily beset by stair- stepping effects, smaller layer thicknesses of 0.127 mm result in a reduction in surface roughness of more than 20% in comparison to the same part when produced with a layer thickness of 0.178 mm.

The Rz value from part sections which are in contact with the support material, in contrast, show no significant effects of varying layer thickness.Surface Finishing through Vibratory GrindingThe term “vibratory grinding” derives from the principles behind the process, in which the grinding action is generated by relative movement, caused by machine vibration, between the part and the abrasive media. In DIN 8580 vibratory finishing processes are categorized with the primary category “Separation Methods” [9]. The implementation of the vibratory grinding principle can entail drum-based tumble grinding, classic vibratory grinding, centrifugal-force- based grinding, or wet grinding. For FDM parts made of Ultem*9085, experiments have already been carried out to which investigate the effects of disc grinding units, which operate on the basis of centrifugal force; increased wear on the edges and corners has been noted when using this option.

[11]The experiments carried out here were done to analyze the effectiveness (= reduction in roughness / finishing time) of a vibratory grinding process on ABS-M30 specimens for various grinding times and grinding intensity levels. The specimens as well as the ceramic abrasive media used are shown in Figure 3 below.c)3d scan ofcorners and edgese)RXX 10/15 Sa)75x50x2mm90 90 45 45 b)25x20x2mmFigure 3:Finishing Media and Specimen DesignThe use of two different-sized specimens (cf. Figure 3 a) and b)) with the same geometry should allow conclusions to be drawn about the influences of part size (particularly in the case of small parts) on the grinding process. The wear on corners and edges due to the grinding process will be analyzed later using three-dimensional measurements of the rectangular specimens (cf. Figure 3 c)).

The wear on corners and edges due to the grinding process will be analyzed later using three-dimensional measurements of the rectangular specimens (cf. Figure 3 c)). Each side of the specimen should be plane; the overhanging faces on the ends of the specimens are intended to prevent the parts from adhering to the wall of the grinding unit. The specimens with pillars of varying lengths and widths (cf. Figure 3 d)) are used to characterize the resistance of the parts against forces to which they are subjected during the grinding process. The grinding media used consist of diagonally cut, triangular-faced rod segments of ceramic; these segments, model type RXX 10/15 S, are from the company RÖSLER Oberflächentechnik GmbH and have face side dimensions of 10x15 mm (cf. Figure 3 e)).In the following, the surface roughness of specimens produced with 90° and 45° build orientations will be examined under different FDM build conditions and for various grinding process parameters.

Due to the build-angle-dependent surface structures resulting from the process itself, the results are continually differentiated based on build orientation. Figure 4 shows the Rz values for specimens built with a 90° build orientation and with a surface area of 75x50 mm, in relation to the layer thickness, grinding time, and grinding intensity.T10 RXX 10/15 S I=80306012001020304050607080finishing time / minRz / µmT12 RXX 10/15 S I=10T10 RXX 10/15 S I=10T12 RXX 10/15 S I=8Figure 4:Specimen Size 75x50x2 mm with 90° Build Orientation: Comparison of Different Layer Thicknesses and Finishing IntensityDue to the lower layer thickness, the curves representing the T10 series show smaller roughness values. The parallel curves in Figure 4 show a constant effectiveness of this vibratory finishing process on parts with different layer thicknesses.

The parallel curves in Figure 4 show a constant effectiveness of this vibratory finishing process on parts with different layer thicknesses. Increasing the grinding intensity, which can be set at levels between 1 and 10 on the trough grinding machine with the help of a potentiometer, appears to result in a slightly better finish.

Increasing the grinding intensity, which can be set at levels between 1 and 10 on the trough grinding machine with the help of a potentiometer, appears to result in a slightly better finish. After a grinding time of 120 minutes on intensity level 10, the average height of the surface roughness profile could be reduced by up to 50%, from 62 µm to 30 µm.The results of a series of tests similar to those described above for specimens with thedimensions 25x20 mm are shown in Figure 5 below.T12 RXX 10/15 S I=8T12 RXX 10/15 S I=10T10 RXX 10/15 S I=10T10 RXX 10/15 S I=80306012001020304050607080finishing time / minRz / µmFigure 5:Specimen Size 25x20x2 mm with 90° Build Orientation: Comparison of Different Layer Thicknesses and Finishing IntensityThe relationship between the intensity level and different layer thicknesses has shown itself to be constant in this case as well.

However, it is noteworthy in comparison to the larger specimens represented in Figure 4 that, here, the finishing process is less effective throughout the entire finishing period. Taking into consideration the somewhat lower initial values, the surface roughness is reduced by only 30%, from 53 µm to 37 µm, for parts with a layer thickness of 0.127 (T10 tip) and an intensity level of 10. One speculative reason for this difference could be the lower resistance of the smaller part to the grinding pressure of the finishing media, because smaller parts tend to move with the flow of the finishing media in the processing bowl.

The influence of the finishing process on the amount of material separated from the specimen can be seen in Figure 6.T10 size 25x20 mmT12 size 25x20 mmT12 size 75x50 mm00,010,020,030,040,050,060,070,083060120decrease of height / mmfinishing time / minT10 size 75x50 mmFigure 6:Material Removal for Specimens Built at 90° Build Orientation Using the Finishing Media RXX 10/15 S at Intensity Level 10The material removal levels shown in the graph above were determined by measuring the specimen thickness with hand-held calipers after varying grinding times. The amount of material removed is the difference between untreated specimens and the specimens which were subject to the finishing process for a defined length of time; due to processing, material was removed from both sides of the finished specimens. As already noted when considering the grinding effectiveness, specimens with larger surface areas appear to exhibit more material removal than those with less surface area.

As already noted when considering the grinding effectiveness, specimens with larger surface areas appear to exhibit more material removal than those with less surface area. Material removal from the larger T12-specimens was ca. 0.07 mm after 120 minutes. In this case, the part measured had a desired thickness of 2.00 mm; directly after production, this value was determined to be 2.07 mm, meaning that with material removal during finishing, the desired thickness was obtained. The T10-specimen with dimensions of 75x50 mm had a post-production thickness of 2.02 mm; in this case, after only 30 minutes of finishing, measurements showed material removal of some 0.04 mm, meaning a deviation downwards from the desired mass to 1.98 mm.

This example serves to demonstrate that the length of the finishing process must be determined in relation to the initial and to the desired dimensions of the finished product in order not to negatively impact dimensional accuracy.Figure 7 shows the change in average surface roughness Rz as a function of grinding timefor parts with layer thicknesses of 0.127 mm and 0.178 mm.T12 size 25x20 mmT12 size 75x50 mmT10 size 75x50 mm03060120020406080100120finishing time / minRz / µmT10 size 25x20 mmFigure 7:Average Height of Profile Rz in Relation to Finishing Time for Specimens with 45° Build Orientation, Using the Finishing Media RXX 10/15 S with Finishing Intensity Level 10 (measurement on top side, perpendicular to the layers)For layer thicknesses of 0.178 mm (T12 tip), there are no significant differences apparent in the results for small and large parts (cf. Figure 7).

Figure 7). In consequence of the pronounced stair- stepping effects affecting parts with a build orientation of 45°, the surface roughness could only be reduced by 20%, even after a 120-minute finishing process, from 118 µm to ca. 90 µm. For 45°-oriented specimens with a layer thickness of 0.127 mm (T10 tip), the initial Rz value (i.e., before finishing) of 88 µm is already below that of T12-specimens after a 2-hour finishing process. In this series of experiments, specimens with larger surface areas also showed themselves more responsive to the finishing process; after a finishing time of 120 minutes, an average Rz value of ca. 49 µm was observed. This translates to a reduction in surface roughness of more than 40% compared to the initial state.

49 µm was observed. This translates to a reduction in surface roughness of more than 40% compared to the initial state. The results of material reduction from the 45°- oriented specimens are shown in Figure 8 below.T12 size 75x50 mmT10 size 75x50 mmT10 size 25x20 mmT12 size 25x20 mm00,010,020,030,040,050,060,070,080,093060120decrease of height / mmfinishing time / minFigure 8:Material Reduction for Specimens with 45° Build Orientation Using the Finishing Media RXX 10/15 S with Finishing Intensity Level 10Figure 8 shows increasing material removal as the length of the finishing process increases. Almost all of the specimens exhibit similar removal rates according to the respective finishing times.

Almost all of the specimens exhibit similar removal rates according to the respective finishing times. The poorer finishing effect seen in Figure 7 for parts with higher layer thicknesses cannot, therefore, be explained by lower levels of material removal; instead, it can be attributed to the more pronounced stair-stepping structure, which creates deeper furrows in the surface of the part produced.A comparison of vibratory finishing and abrasive blasting with glass spheres (diameter70-110µm) is presented in Figure 9.0.050-50[µm]Abrasive Blasting Spheres 70-110µm 20 secVibratory FinishingRXX 10/15 S 120 minFigure 9:Surface Roughness for Parts With 90° Build Orientation (measurement perpendicular to the layers on top side, size 75x50x2 mm, T12 tip)After a finishing time of only 20 seconds (abrasive blasting), the measured Rz value (51µm±4µm) is similar to the value after 120 minutes of mass finishing (51µm±4µm).

At first this seems to be a promising result, but after analyzing the surface profiles after finishing (cf. Figure 9), it was observed that the blasting process not only grinds the part surface, but also removes material within surface troughs, resulting in no absolute change in overall roughness profile height. The vibratory finishing process smooths only the top of the filaments; therefore the results of abrasive blasting have to be taken with a grain of salt when focusing on smooth part surfaces.Due to the nature of the process, vibratory grinding leads to mechanical stress on the parts. The analysis of FDM tensile testing specimens made of Ultem*9085 which had been finished on a disc grinder was not able, despite testing some specimens to destruction, to offer reliable conclusions about the influence of the finishing process on the characteristic mechanical values.

Therefore, in order to attempt to characterize mechanical loading during the finishing process, specimens were created specifically for the test which featured several pillar-structures of varying length and with varying cross-sectional areas. The specimens were oriented in the build chamber such that these pillars were produced in the critical upright build direction (the Z- direction). [11]lAFigure 10:Analysis of Part Destruction after Grinding Pillar-SpecimensBased on the varying cross-sections and lengths of the pillars, the number of pillars broken off after a finishing process offers one possibility of drawing conclusions about the intensity of the mechanical loading for different grinding media, intensity levels, or grinding times. From a mechanical perspective, the pillars, as attached to the base, can be seen as cantilever beams.

From a mechanical perspective, the pillars, as attached to the base, can be seen as cantilever beams. In a collision with the grinding unit components or other elements, the force applied to the pillar tip can be simplified as a bending moment in the support. Using the bending moment equation for a simple cantilever beam with a rectangular cross-section, a stress factor of A²/l³ can be determined with the following relationship: [11] [12]EclAB432(1)Here, A is the cross-sectional area (mm2), l the length of the pillar (mm), E the Young’s Modulus (N/mm2), and cB (N/mm) the spring constant of the cantilever.

Figure 11 shows the percentage of intact pillars for 4 specimens after finishing in terms of this stress factor.intensity: level 8intensity: level 100204060801000,0000,0200,0400,0600,0800,100intact specimens / %stress factor A²/l³ / mmFigure 11:Intact Specimens Built with a T10 Tip after 120 Minutes Finishing Time with Grinding Media RXX 10/15 SFor a stress factor A²/l³ > 0.016 mm (meaning larger cross-sections in relation to length), there was no destruction registered for any of the specimens finished at intensity level 8. For higher grinding intensity (Level 10), part damage could be observed in the specimens starting at astress factor of 0.032 mm.

For higher grinding intensity (Level 10), part damage could be observed in the specimens starting at astress factor of 0.032 mm. Thus, when using the abrasive media RXX 10/15 S, higher grinding intensity has a negative effect on part stability.The influence of grinding time and intensity on the dimensional exactness as registered bythree-dimensional measurements on a rectangular specimen is shown in Figure 12.Intensity: level 10Intensity: level 8Finishing time 30 min120 min0.200.100.00-0.10-0.20Erosion / mmFigure 12:3D Scan of Specimens Built with T10 Tip, Finished with RXX 10/15 SThe color changes seen on the specimen scans in Figure 12 represent local geometrical deviations from the desired dimensions in comparison to unfinished specimens. No measureable difference between intensity levels was observed here on the large surfaces when finishing parts with a layer thickness of 0.127 mm while using RXX 10/15 S.

No measureable difference between intensity levels was observed here on the large surfaces when finishing parts with a layer thickness of 0.127 mm while using RXX 10/15 S. An examination of specimens with a finishing time of 120 minutes reveals increased material removal from the long edges related to higher intensity. At a mere -0.03 mm, however, this difference is very slight. At the corners of the specimens, scattered chamfering and rounding up to 0.1 mm was observed for finishing times of 120 minutes, which is significantly higher than average material removal levels. When using larger abrasive media, similar abrasive effects can be observed; however, corners and edges are much more rounded in this case.SummaryAn analysis of the surface structure of unfinished FDM parts made of ABS-M30 revealed smaller differences in surface roughness than initially expected for differing layer thicknesses of 0.127 mm (T10) and 0.178 mm (T12 tip).

In particular for areas where the part is in contact with the support material, the process used exhibited no measureable difference between different layer thicknesses. The difference in resolution is, however, of great importance for part surfaces produced with a build orientation of 45°; here, an increase in average surface roughness Rz of more than 20% could be observed. Vibratory grinding has shown itself to be an especially mild finishing process with regard to rounding of edges and part damage. Lower grinding intensity isespecially advantageous in order to minimize edge and corner rounding. When using the ceramic grinding media RXX 10/15 S, lower intensity levels result in a reduction in grinding effectiveness compared to the highest intensity level; this reduction is, however, slight.

When using the ceramic grinding media RXX 10/15 S, lower intensity levels result in a reduction in grinding effectiveness compared to the highest intensity level; this reduction is, however, slight. Thus, using lower grinding intensity not only offers advantages with respect to dimensional accuracy, but is no less advisable in terms of grinding effectiveness.AcknowledgementsThe authors want to thank all industry partners of the DMRC as well as the federal state of North Rhine-Westphalia and the University of Paderborn for the financial and operational support within the project “Analysis of the FDM Part Quality manufactured with ABS M30 with the focus on the toy industry”.References[1] T. Wohlers: „Wohlers Report 2013 – Additive Manufacturing and 3D Printing State of the Industry“, Annual Worldwide Progress Report, 2013 [2] A. Gebhart: „Generative Fertigungsverfahren / Rapid Prototyping – Rapid Tooling – Rapid Manufacturing“; Carl Hanser Publisher, München 2007 [3] I.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFinishing of Additively Manufactured Metal Parts by Abrasive Flow MachiningXuanping Wang1, ShichongLi1, Youzhi Fu1, Hang Gao1,*1School of Mechanical Engineering, Dalian University of Technology, Liaoning, China,116023Corresponding Email: gaohang@dlut.edu.cnAbstractSurface finishing is still a crucial challenge in metal Additive Manufacturing (AM) as the as-built surface roughness is difficult to fulfill service requirements, due to staircase effect, balling effect inherent to AM. Abrasive flow machining (AFM) is a non-conventional finishing technique that offers better accuracy and efficiency for parts with difficult-to-access structures, and the application of AFM to finishing metal parts of AM process is discussed in this paper.

The aluminum and titanium grilles by selective laser melting are taken to explore the finishing effect of outer and inner surfaces. The AFM process parameters of abrasive grits sizes, abrasive media viscosity, and tooling designs are optimized to implement effective material removal from the outer and inner surfaces. The results show that the AM grille parts with non-trial internal structures can be finished efficiently and consistently by AFM.Key words: additive manufacturing, abrasive flow machining, polishing, complex structureIntroductionThe surface roughness of parts additively manufactured by SLM is able to reach Ra 10μ m ~50μm, while the surface roughness of parts by subtractive manufacturing is usually below Ra 2.5μm. As the inherent "balling effect","powder adhesion" in additive manufacturing lead to the poor surface roughness[1-3], polishing process is inevitable to fulfill requirement of surface quality.Although blasting, belt polishing, laser polishing, et al.

are applied in polishing metal AM parts, it is still challenging to polish AM parts with difficult-to-access complex geometries[4]. Abrasive flow machining (AFM) is a polishing process with semisolid abrasive media as cutting tool. The deformable characteristics of the abrasive media enables it to act as a flexible cutting tool to achieve high surface quality for parts with complex geometries[5]. In this paper AFM is taken to polish additively manufactured grille to explore the finishing effect of outer and inner surfaces.Experimental setupAs shown in Fig.1, the grille is made by aluminium alloy powder through Selective Laser Melting(SLM), on which inclined square holes with 3mm long each side are distributed in equal distance. As shown in Fig.2, there are obviously alloy balls from the SLM process and alloy powders adhered on the grille surface, which results in the poor surface quality with surface roughness Ra 14μm.2470(a) Outer surface (b) Inner surfaceFig.1.

Initial surface of the grille Fig.2. Initial surface textures of the grilleIn AFM polishing process, abrasive media act as the continuously deforming cutting tool, composed of polymer, abrasive grits, et al. When abrasive media are extruded across the workpiece surface, the active abrasive grits scratch the surface and micro-chips are produced, leading to the material removal. The AFM machine used in the present work is shown in Fig. 3, composed of hydraulic system, electric system, and mechanical system. As the tooling and the flow channel demonstrated in Fig.4, it is optimized to implement polishing the outer surface and inner surface of the grille simultaneously.Fig.3. AFM systemFig.4. tooling and flow channel for grille polishingExperimental results analysisMesh size and mass fraction of the abrasive media, and the process parameters in AFMpolishing are listed in Table 1.Media SiC24#-80#Table 1.

Processing parameters in AFM polishing Time /min Cycles 23 50Pressure /MPa 2.5Angle /° 2After AFM polishing, the surface roughness is reduced from the initial Ra 14μm to Ra0.94μm. According to the micro-profiles of outer surface and inner surface provided in Fig.5, the adhered melten balls and alloy powders are all removed with comparison to the initial surface textures in Fig. 2.2471It could be concluded from the experimental results that AM parts with non-trivialinternal structures can be polished effectively, including outer surface and inner surface.(a) Outer surface (b) Inner surface Fig.5. surface textures of the grille after AFM polishingConclusionsIn this paper abrasive flow machining(AFM) technique is taken to polish additively manufactured grille with square inclined holes with regular arrangement. Abrasive media with hybrid grit sizes are adopted to implement polishing the inner and outer surface of the grille simultaneously.

Abrasive media with hybrid grit sizes are adopted to implement polishing the inner and outer surface of the grille simultaneously. After AFM polishing process, the poor surface resulting from the inherent "balling effect", "powder adhesion" in additive manufacturing is significantly improved, proving that AFM is a feasible way to polish additively manufactured parts with difficult-to-access complex geometries.References [1] Frazier W E. Metal additive manufacturing: A review. Journal of Materials Engineering and Performance, 2014, 23(6): 1917-1928. [2] Lamikiz A, Sanchez J A, de Lacalle L N L, et al. Laser polishing of parts built up by selective laser sintering. International Journal of Machine Tools & Manufacture, 2007, 47(12-13): 2040- 2050. [3] Rosa B, Mognol P,Hascoet J Y. Laser polishing of additive laser manufacturing surfaces. Journal of Laser Applications, 2015, 27: S29102-1. [4] Strano G, Hao L, Everson R M, et al. Surface roughness analysis in selective laser melting.

NISHING OF SLS-PARTS FOR RAPID MANUFACTURING (RM) – A COMPREHENSIVE APPROACH –M. Schmid*, C. Simon* and G.N. Levy*INSPIRE, irpd – Institute for Rapid Product Development, Lerchenfeldstr. 5, CH-9000 St. Gallen, Switzerland Reviewed, accepted September 15, 2009AbstractPlastic parts are often coated to fulfill the desired functional requirements during product life. This may be for decorative purposes only, but also for functions such as improved tribology, wear and humidity resistance, UV- and light stability, hygienic and biofilm resistance. Moving SLS towards Rapid Manufacturing (RM) and making those parts competitive with parts produced by other techniques (e.g. injection molding) implies the adoption of a new quality of part finishing and coating strategy for SLS. This paper provides a survey of current finishing methods used for RM-SLS parts in our institute and highlights the manually-driven process steps.

This paper provides a survey of current finishing methods used for RM-SLS parts in our institute and highlights the manually-driven process steps. The need for, and first trials with, a more automated finishing process (e.g. vibratory grinding) are discussed, as is the need for innovative supporting software tools.IntroductionOver the last two decades more than 30 different layer manufacturing (LM) technologies or “additive processing technologies” based on different technological principles have been developed. The most promising technologies have been engineered to a higher level and are now used for the production of prototypes, parts or models in a great number of applications [1-3]. LM technologies are divided into rapid prototyping (RP) and rapid manufacturing (RM) technologies, depending on the end use. RP involves the production of prototypes, visual design aids and test parts, while RM involves the production of actual production parts (end products).

RP involves the production of prototypes, visual design aids and test parts, while RM involves the production of actual production parts (end products). At present, the market for RM applications is experiencing rapid growth [4], but suffersfrom certain drawbacks that hinder greater acceptance in the industry. Apart from the lack of several suitable polymeric materials for layer wise production [5], In general, the surface quality of SLS parts is relatively poor for various reasons [6]. Especially when SLS parts are compared with plastic parts produced with “classical” injection molding (IM) technology, there is an obvious difference in surface quality, as IM uses pressure and polished tools. Since, in our modern world, most plastic parts are coated for various reasons, RM has to take account of this market reality and develop strategies in order to remain competitive. Otherwise, the great advantage of SLS of increased complexity is obsolete due to inferior surface quality.

Otherwise, the great advantage of SLS of increased complexity is obsolete due to inferior surface quality. This means that RM has to develop some basic methods for improving surface roughness in particular. This cannot be a manually-driven process as in RP (e.g. priming and polishing with abrasive papers), as a lot of parts must achieve the same quality that has to be guaranteed to the customer in advance. Previously described laser-based finishing approaches [8, 9] are also unusable due to the higher part numbers. One promising method that may satisfy this requirement for reproducible surface refining is vibratory grinding. Furthermore, functional coatings for SLS parts must be systematically evaluated in future, and their effectiveness for certain applications must also be demonstrated to customers in advance in order to fulfill the required specifications defined forintended applications.

This paper describes the first trials for vibratory grinding and some successful results in terms of better surfaces, and highlights the urgent need for software improvements in order to help protect SLS parts on exposed sites. Furthermore, new functional coatings will be introduced and measurements carried out to prove their desired functionality. There has been little systematic research to date on this topic [10, 11].Experimental Part (Methods)Vibratory grinding: The vibratory grinding experiments were performed on a rotary vibrator manufactured by Roesler: R180/530 TE-30 RMO.

Ceramic media with the following identification (type/size (mm)/geometry) were used as abrasive grinding bodies, as recommended by Roesler: RXX 10/12 DZS, RS 03/10 ZS, RXX 04/04 S (RXX = highest possible abrasion; RS = medium abrasion; DZS = trident diagonal cut; ZS = cylinder diagonal cut; S = triangle diagonal cut); medium: water; frequency of vibration: 50 Hz; rotation: 3000 rpm in general for all trials.

Roughness analysis: The roughness analysis was performed with a surface measuring instrument: Perthen Profilometer (Perthometer S3P with RFHTB-250 sensor); measuring distance: 5.6 mm; tactile force: 0.6 mN; the measurement was reproduced 5 times for each sample and a median value was calculated; Ra (arithmetic average of absolute values) was taken as the main evaluation parameter; Optical analysis: The optical analysis for comparison of parts before and after vibratory grinding was performed with Atos II, System Configuration: measured points: 1,400,000; measurement time: 1 second; point spacing: 0.12 - 1.4 mm; Dip coating: Coating was done with test plates suitable for hydrostatic head testing (diameter: 65 mm, thickness 2 mm); all test plates were immersed for 10 minutes in the coating solution, dried after removal overnight under vacuum and repeated (double coating); all coating solutions had a concentration of 10% (w/w) in an organic solvent according to the supplier's recommendation (water, ethanol or acetone).

The respective active component was derived from the group of silicons, acrylates, vinyl-acrylates, polyurethanes or epoxides; Infrared spectrometry: Infrared spectrometry for the qualitative analysis of coatings was performed with a BIORAD 75 FTS spectrometer equipped with an HATR measurement chamber for surface-sensitive measurements (horizontal reflection); no sample preparation was required; Water tightness: The tightness of the samples regarding water permeation was analyzed with a hydrostatic head tester usually used for textiles (Company Textest; Typ: FX 3000); the maximum water pressure is equivalent to 1 bar or a 10 m water column; the stop criterion for the measurement was the appearance of three water droplets on the surface (standardized for textiles ISO 13859); Water repellence: Contact angle measurements were performed with a System G 10 (Krüss GmbH) with deionized water; measurements were repeated 5 times to obtain a median value; Wear resistance: The abrasion resistance was measured with a Taber-Abraser testing device by weight loss (gravimetric) after 100 and 500 rotations; friction roller type: H18; load: 500 g on every roller; due to the non standardized roller type for plastics, the measurements could not be treated as absolute values but can be compared among themselves;Results and Discussion, Part 1: Surface RefiningAs long as SLS parts are intended to be prototypes a hand finish is acceptable since there is usually no preliminary defined quality control, desired surface or roughness.

Hand finishing inthis case involves a process used in many other connections (e.g. car repair) and consists mainly of three steps: application of a primer, polishing of surfaces and cosmetic spraying. Figure 1 shows the steps required, taking the example of a part (industry cover) produced in our factory. It is obvious that this process is - especially for the polishing step (red circle) - labour-intensive and requires specialized staff with certain skills (e.g. car body painter). Furthermore, the process is time-consuming and expensive.Primer PrimerPolish PolishSpray Sprayvarnishable varnishablepriming primingpolishing by abrasives polishing by abrasivestraditional look traditional lookFigure 1: Composition of the process steps of surface refining with manual polishing;To overcome the drawbacks of the manual process we investigated, in collaboration with the company Roesler, the possibility of vibratory grinding for SLS parts and the corresponding effects.

Please note that vibratory grinding is an entirely empirical process and can be validated only by experiment. Several test parts with a variety of selected geometries (planar surfaces, parts with edges and curving geometries) were chosen in order to obtain comprehensive information on achievable surface quality. The main factors influencing the results would be expected to be the ceramic media (abrasion body), the rotation speed and the time of exposure. Figure 2 gives an overview of the selected abrasion bodies. As regards the nomenclature of Roesler, RXX media are those with the highest abrasion performance available. RS media produce medium abrasion. The geometrical size of the media was selected to match the size of the SLS parts.RXX 10/12 DZS RXX 10/12 DZSRS 03/10 ZS RS 03/10 ZSRXX 04/04 S RXX 04/04 SFigure 2: Assembly of the ceramic media employed for the vibratory grinding testsFigure 3 summarizes the parts and some of the qualitative results.

These qualitative appraisals were produced after visual inspection and represent subjective impressions. In this context a running time of 4 hours is not sufficient (test 1). A grinding period of at least 12 hours (test 3, 4) or 8 hours (test 5, 6) is required. The best results for surface quality for all cases were found for test 5, which produced a homogenous and acceptable surface. Due to the geometrical conditions of some test parts, the abrasion bodies became jammed within these parts. This undesirable effect was found for test 5 and other tests (see Figure 3: ‘chips became blocked’).Test Nr.

Test Nr.Parts PartsChips ChipsDescription of results Description of results1 (DF) 1 (DF)2 (DF) 2 (DF)RXX 10/12 DZS RXX 10/12 DZSRXX 10/12 DZS RXX 10/12 DZSAfter 4 hr, surface removal was After 4 hr, surface removal was inadequate inadequate As for test 1 but after 8 hours; the As for test 1 but after 8 hours; the surfaces were better, but still not surfaces were better, but still not clean enough clean enough3 (DF) 3 (DF)RXX 10/12 DZS RXX 10/12 DZSThe surface was well polished The surface was well polished after 12 hours; several parts were after 12 hours; several parts were damaged damaged4 (DF) 4 (DF)RS 03/10 ZS RS 03/10 ZSThe removal rates for the RS chips The removal rates for the RS chips were inadequate after 12 hours were inadequate after 12 hours5 (HST) 5 (HST)RXX 04/04 S RXX 04/04 SThe surfaces of these parts were The surfaces of these parts were homogeneous and acceptable; homogeneous and acceptable; some chips became blocked, some chips became blocked, duration 8 hours duration 8 hours6 (DF) 6 (DF)RXX 04/04 S RXX 04/04 SThe efficiency for the removal of The efficiency for the removal of small chips on large parts was small chips on large parts was inadequate; the surfaces were inadequate; the surfaces were inhomogeneous and some chips inhomogeneous and some chips became blocked became blockedFigure 3: Compilation of test parts for vibratory grinding and qualitative appraisalIn conclusion, these results give a preliminary indication about suitable grinding bodies (ceramic media) and the minimum process time needed to achieve a positive result for the investigated material.

Further investigations are required for other SLS materials.More precise measurements of the surface roughness of the parts by profilometer after the different tests are summarized in Figure 4 for the purposes of comparison. The measured surface parameter Ra (arithmetic average of absolute values) shows, in a very distinct manner, that the surface quality of all parts is significantly improved compared to a “regular” part made from the same material by SLS (sample DF virgin). However, despite the visual examination in Figure 3, the surface quality in tests 1 to 5 did not differ as greatly as might be expected from the qualitative evaluation of profilometer measurements. All surface roughnesses (Ra) are around 2 µm and equal within the standard deviation. However, the standard deviation for test 1 and test 2 is significantly greater compared to the other tests, and test 5 tends to produce the smoothest surface.

However, the standard deviation for test 1 and test 2 is significantly greater compared to the other tests, and test 5 tends to produce the smoothest surface. Accordingly, this result is also in accordance with the qualitative appraisal.14.0Ra12.010.08.06.04.02.00.0DF virgintest 1test 2test 3test 4test 5Figure 4: Surface roughness (Ra) before and after vibratory grinding (DF virgin = untreated part)In addition to this qualitative and profilometer analysis, a space-resolved optical analysis (white light scanner with grid projection) was performed after vibratory grinding to obtain information about the material layer thickness abrasion. The result is a variance comparison between the SLS sintered part and the same part after vibratory processing. Figure 5 shows two parts that were analyzed by this technique. The colors represent the abrasion rate at different points.

Figure 5 shows two parts that were analyzed by this technique. The colors represent the abrasion rate at different points. Blue represents no, or minimal, abrasion, while red indicates maximum abrasion loss of up to 250 µm. We can estimate a typical abrasion profile < 0.1 mm, but the maximum may even exceed 0.3 mm (indicating part damage).Part 2 Part 2Part 1 Part 1Figure 5: Optical analysis of parts after vibratory grinding and typical profileIt is readily apparent that the results are as expected. Flat surfaces (Part 1) exhibit a very homogenous appearance with the desired optimized surface. As the edges are curved no dramatic weight loss is detected here either. A different result can be seen for Part 2 in Figure 5. The edges are severely damaged (red regions = abrasion loss of more than 250 µm).This means that further basic research is required here.

A different result can be seen for Part 2 in Figure 5. The edges are severely damaged (red regions = abrasion loss of more than 250 µm).This means that further basic research is required here. The main targeted research approach at the moment therefore focuses on the development of new software capable of detecting problematic regions in part geometry semi-automatically, so that these can then be covered with protecting aids. These auxiliary constructions must be arranged in such a way that they can easily be removed after processing. Further work should focus on a family-dependent media definition (abrasion media), optimized process parameters and suitable quality control after vibratory grinding.Results and Discussion, Part 2: Functional CoatingIn addition to the surface optimization (reduction of surface roughness (see Part1)), further enhancement is also needed for the application of functional coatings for RM in order to improve the parts for desired applications.

This can be e.g. improved water tightness, humidity resistance, wear advancement, UV- and light stability, biofilm resistance, etc. Furthermore, a functional coating can sometimes help to compensate for the lack of feasible material grades for SLS.For this reason a dip coating (DC) process was established in order to facilitate the application of different coatings. Figure 6 illustrates the dip coating procedure. Basically, a DC solution is provided by dispersing an active chemical component with the desired functionality in an organic solvent or water. The part is then simply immersed in and removed from this solution. The great advantage of DC is the acceptance of any complexity of part geometry and the cheap and simple approach. Furthermore, the thickness of the layer can be controlled to a certain extent by modifying the concentration of active component in the solution, dwell time or removal action.

Furthermore, the thickness of the layer can be controlled to a certain extent by modifying the concentration of active component in the solution, dwell time or removal action. Disadvantages of DC include the large amount of solution required for the complete immersion of larger parts and, in some cases, limited stability of the coating chemicals dissolved in the DC solution.dip coating process dip coating processImmersion ImmersionDwell time Dwell timeWithdrawal WithdrawalFigure 6: Functional layers applied by the dip coating (DC) processVarious chemicals recommended by distributors for this purpose were utilized for the test that was basically designed to improve the water tightness of SLS parts. These chemicals belong to the group of silicones, acrylates, vinyl-acrylates, polyurethanes, etc. and were applied by DC to a disk specimen with a diameter of 60 mm and a thickness of 2 mm.

These chemicals belong to the group of silicones, acrylates, vinyl-acrylates, polyurethanes, etc. and were applied by DC to a disk specimen with a diameter of 60 mm and a thickness of 2 mm. The presence of a coating layer on the specimen was checked qualitatively by surface sensitive infrared spectrometry (HATR-IR).A hydrostatic water column tester (Textest FX 3000) was used for measuring watertightness. This measuring equipment is usually used for measuring the effectiveness of water- tight textiles, but was also found to be suitable for the assessment of coatings on SLS parts. For the test the coated SLS disks were sealed over the water bath of the FX 3000, and the pressure in the water was then increased linearly. As defined in textile standards (e.g. DIN 53886) the test is concluded when 3 droplets of water are visible on the dry side of the specimen. Figure 7 summarizes the results of the water tightness test for all applied coatings after a first and a second DC procedure.

Figure 7 summarizes the results of the water tightness test for all applied coatings after a first and a second DC procedure. The water pressure [mbar] measured when three droplets are visible is noted.100> 1000> 1000one time DC two times DC] r a b my l e t a i d e m my l e t a i d e m m[e r u s s e r p r e t a w50i d e b r o s d ar e t a wi d e b r o s d ar e t a w02 12 1 A P1 16 13 243 27 28 22 33 5 7 31 29 25 20 1(material code numbers)2 22 136 15 11 4 38 19 16 2SiliconePolyurethane Vinyl-Acrylpolymers DiversFigure 7: Results of water tightness measurements of different coatings on the SLS disk specimen in comparison with uncoated PA12 (Duraform)The following conclusions can be drawn from this figure. Certain coating chemicals improve the water tightness of PA12 (uncoated Duraform) as desired, whereas others even reduced water tightness.

Certain coating chemicals improve the water tightness of PA12 (uncoated Duraform) as desired, whereas others even reduced water tightness. Polyurethane (PU) seems to be an unsuitable coating for water tightness purposes, although PU is very often used in practice. Silicones and vinyl-acrylates showed the desired positive effect of improving water head pressure to PA12. Whereas silicones need just one DC step and are not significantly improved by a second cycle, acrylic polymer seals need two or,0 3most likely, even more DC runs. Two coatings, in particular, are outstanding at both ends of the scale. Numbers 11 (silicon) and 15 (vinyl-acrylate) shows phenomenal water tightness in both cases for the second coating, until the maximum measuring range of the equipment is reached. The other end of the scale is represented by coatings 25 and 26. In these cases, the water droplets appeared on the surface immediately after the specimen was sealed.

The other end of the scale is represented by coatings 25 and 26. In these cases, the water droplets appeared on the surface immediately after the specimen was sealed. This behavior could even be beneficial if a high water absorbing material is desired for a certain application.Another interesting aspect in this connection is the change in surface energy orhydrophobicity of a specimen with the different coatings. This effect is analyzed by contact angle measurements. The results of these measurements are summarized in Figure 8. If the contact angle is below 90° the surface is classified as hydrophilic, which is the case for almost all polyurethane and acryl polymer coatings. Silicones have a low affinity (surface energy) for water droplets and thus produce predominantly hydrophobic surfaces. A super-hydrophobic surface was achieved for coating number 16 (lotus effect).

Silicones have a low affinity (surface energy) for water droplets and thus produce predominantly hydrophobic surfaces. A super-hydrophobic surface was achieved for coating number 16 (lotus effect). .180super hydrophopicone time DCtwo times DChydrophopic90hydrophilicy e t a d e m mliy e t a d e m mliid e d a e r p s r e t a wd e d a e r p s r e t a w02 12 1 A P1 16 13 243 27 28 22 33 57 31 29 25 20 1(material code numbers)2 22 136 15 11 4 38 19 16 2SiliconePolyurethane Vinyl-Acrylpolymers DiversFigure 8: Contact angle measurements on the different surfaces after coatingAlthough it did not form the focus of this investigation of water tightness improvement, we also analyzed the friction behavior of some of the coatings with the Taber abrasion tester. Figure 9 clearly shows that some selected coatings also improve the wear behavior of PA 12 (Duraform = DF).

Figure 9 clearly shows that some selected coatings also improve the wear behavior of PA 12 (Duraform = DF). Especially coating 16, which exhibited the highest hydrophobicity, also shows significantly reduced abrasion. Further research regarding this topic will be necessary and is already anticipated in our group.0 345.0100-500 ro tatio ns40.035.00-100 ro tatio ns] g m[s s o lt h g i e w30.025.020.015.03 . 7 17 . 3 17 . 77 . 81 . 32 . 4 110.05.04 . 7 17 . 89 . 1 19 . 4 17 . 2 10 . 3 10.0DF1115 13 s amp le c o d es1633Figure 9: Abrasion behavior of different coatings compared to virgin PA 12;Conclusion and OutlookThis paper shows that vibratory grinding is a promising approach for refining SLS parts from the standpoint of surfaces. In progressing from RP to RM this is an essential step in guaranteeing the surface qualities and process reliability required by customers.

In progressing from RP to RM this is an essential step in guaranteeing the surface qualities and process reliability required by customers. Suitable ceramic bodies were investigated and process times were determined. In the optical variance analysis we demonstrated that parts without complex structures are readily amenable to vibratory grinding (removal < 0.1 mm). Further research is necessary to improve the protection of the more critical parts for the process in its sensitive regions, and process parameters for vibratory grinding also need to be refined.The application of functional layers to SLS parts by dip coating was shown to be a simple and fairly powerful tool in this work. Chemical coatings that improve the water tightness of PA12 parts were successfully investigated. Furthermore, the change in surface energy and abrasion resistance due to the surface coatings was also demonstrated.

Furthermore, the change in surface energy and abrasion resistance due to the surface coatings was also demonstrated. Moreover, the method is also to be used in future for investigations on UV-protecting coatings and flame retardancy. Additionally, functional coatings may sometimes help to compensate for the lack of material grades that are suitable for SLS.AcknowledgementsI am grateful to the whole laboratory staff of our Institute who supported this work, especially Martin Eichenberger who applied all the coatings and carried out almost all the analytical work.Literature1 Kruth J.-P., Levy G. et al., Consolidation phenomena in laser and powder-bed based layeredmanufacturing CIRP Annals - Manufacturing Technology, 56(2), 2007, 730-759 2 Levy G., Schindel R. et al., Rapid Manufacturing and Rapid Tooling with LayerManufacturing (LM) Technologies, State of the Art and Future Perspectives, CIRP Annals - Manufacturing Technology, 52(2), 2003, 589-6093 Wendel B., Rietzel D.

E Reeves, R C Cobb Centre for Rapid Prototyping Manufacturing Department ofManufacturing Engineering and Operations Management University ofNottingham University Park, Nottingham NG7UKABSTRACTThe use of StereoLithography (SL) can produce accurate prototype models with complex internal and external features. However, a major problem to commercial use is the poor surface finish caused mainly by 'stair stepping' which is inherent in layer manufacturing.

However, a major problem to commercial use is the poor surface finish caused mainly by 'stair stepping' which is inherent in layer manufacturing. Models are often finished by hand but this is labour intensive, highly selective and causes inaccuracies in the model geometry.A three-year research project has been undertaken to address these issues and to investigate This papera range of surface coatings and mechanical finishing processes applied to SL describes some initial findings using resin coatings applied to both cured andSL parts.Initial findings suggest that excess resin retained after part stripping cansurface roughness than parts thoroughly cleaned prior to post curing. photocurable and epoxy based resins to parts in both the un-cured, green and cured roughness has been seen to be reduced by up to 50% on complex parts.a lower addition of surfaceINTRODUCTIONRapid Prototyping can produce parts faster and more economically than by conventional with techniques.

The processes are best suited to parts which are generally complex in freeform curves and re-entrant features, possessing only a limited percentage of plane ......... AJL"""...' ' ' ' ' ......

AJL"""...' ' ' ' ' ...... However, a major problem to commercial use is the poor surface finish caused mainly by stepping' which is inherent in layer manufacturingWhen finishing StereoLithography models best results are obtained following multiplehand finishing in the cured state, with a series of grades of abrasive papersl techniques are highly selective, with finishing of fine detail and internal features many cases finishing of models is neglected as it is labour intensive and not cost """ff.~t"'1"·nT""".surface finishing96With the introduction of new resins and build styles, StereoLithography part accuracy and surface finish has increased almost six fold in as many years, with the ACESTM build style resulting in parts with a roughness average (Ra) of less than 0.28 Jlm on up-facing planes2 • The introduction ofthe new Quick-cast™ 1.1 build style makes possible roughness values of 0.25 - 0.5Jlm Ra for parts used as sacrificial patterns for investment casting3.

Quick-cast™ 1.1 uses triple up-facing layers to reduce deviation by applYing additional layers of material to the part, covering any deviation caused by previous layers. However, research into these new build styles has addressed only the problems associated with vertical and horizontal planes. Analysis of the overall surface roughness of SL parts suggests that the surface roughness on angled planes requires fundamental research (Figure 1).Work has been undertaken at the CAD stage to reduce surface roughness on complex planes and features using part orientation software prior to slicing4 . The CAD representation is oriented about its axis in such a way that stepping will be reduced on planes perpendicular to the 'Z' axis, hence reducing the 'stair stepping' effect. However, the effects of trapped volumes, delamination, build time and the working envelope must also be considered and can prevent build of a part in an optimum orientation.

However, the effects of trapped volumes, delamination, build time and the working envelope must also be considered and can prevent build of a part in an optimum orientation. With many complex parts, software orientation can only reduce 'stair stepping' on a limited number of surfaces and additional post process finishing is still required.research has been undertaken to establish suitable finishing techniques for StereoLithography models using a number of conventional mass finishing systems. The INSTANTCAM projectS during 1992 addressed the problem of SL surface roughness using tumble peening and sandblasting in addition to spray painting. Work at the University of Nottingham applied SL components to conventional mass finishing equipment including barrel tumbling, vibratory finishing, ultrasonic abrasion and blasting6 . The resulting acrylic components were found to have excessive damage, with loss of material at both edges and corners (Figure 2).

The resulting acrylic components were found to have excessive damage, with loss of material at both edges and corners (Figure 2). In both work programs and in work undertaken at the GINTIC institute7 on SL abrasion, the abrasive systems employed were developed for the mass finishing of metallic components using harsh ceramic media.RecentRESEARCH ACTIVITIESGiven the need for an economic method for StereoLithography part finishing, the University ofNottingham has undertaken a three year research project to define a non selective system for the surface finishing of SL models, with minimal manual intervention and loss to the intended part geometry.

A three phase approach to surface finishing has been defined (Figure 3), dependent on the part geometry.Chemical machining of SL parts using acids and solvents to remove excess material from steps defined by the CAD geometry.Abrasive finishing systems including barrel, vibratory and ultrasonic machining using specially selected finishing media to remove material from the steps defined by the CAD geometry.97Additive processing of SL parts applying coating materials to reduce surface deviation by filling the steps within the CAD geometry.Some ofthe initial findings using coating materials for additive processing will be reported inthis paper.order to define the most appropriate processing techniques for the models an understanding of the factors affecting the resolution of SL parts is required.

The factors which affect the overall resolution of a part include resin type, resin temperature and viscosity, part stripping residue, build orientation, CAD and STL representation and the parabolic curing profile of the SL resin in addition to the defined layer thickness. Layer thickness or step height is the major contributory factor in the generation of surface roughness but present resin and machine technology prevents proctical reduction in layer thickness, in addition to the increased build time associated with thinner layers.Research into surface roughness suggests that in addition to layer thickness, excess resin within the build step will result in a change to the surface profile, and hence the surface roughness. The effects of this excess material or 'part stripping residue' can be assessed if a series of part surfaces are compared to the surface defined by the CAD data.

The effects of this excess material or 'part stripping residue' can be assessed if a series of part surfaces are compared to the surface defined by the CAD data. To do this a mathematical representation of a layer manufactured surface can be derived assuming that the layer profile is at right angles, as opposed to parabolic (Figure 4). From this mathematical representation, it is possible through the measurement of a series of SL samples to plot both the mathematical and 'actual' roughness average values for a given layer thickness (Figure 5).A series oftest samples were constructed with incline planes ranging from 10 to 90 degrees, in 10 degree increments, with a measurement of surface roughness average taken at 10 points on the plane surface (Figure 6). As it can be seen the actual surface roughness measured across a range of surface planes give lower Ra values than mathematically predicted.

As it can be seen the actual surface roughness measured across a range of surface planes give lower Ra values than mathematically predicted. The effects of the parabolic curing profile should in theory increase the surface roughness however, analysis suggests that excess resin not removed during part stripping results in a reduction in surface Ra (Figure 7).PHOTOCURABLE RESIN COATINGSAdditional material within the build step can therefore be considered to result in SL parts with lower Ra values than those thoroughly cleaned prior to post curing. This has directed research towards establishing a series of coating materials to further reduce the surface roughness values of SL parts. StereoLithography resin has already been seen to reduce Ra values when not fully stripped during the build step. Initial experiments were thus conducted on both acrylic and epoxy parts by applYing SL resin as a coating to both cured and un-cured 'green' state models.

Initial experiments were thus conducted on both acrylic and epoxy parts by applYing SL resin as a coating to both cured and un-cured 'green' state models. This resulted in parts with excessive loss ofgeometry, due to the high viscosity of the resin coatings. A thinning agent was therefore used to reduce the working viscosity of SL resin. At the lower viscosity, thinner coatings could be applied to parts without excessive loss to part geometry. Low viscosity coatings could then be built up until sufficient material remains in the build step to reduce the overall part Ra, this being a function of the layer thickness and the coating viscosity (Figure 7).98use of StereoLithography resin as an additive coating will result in parts which are both mechanically and visually as near to 'virgin' SL part as possible. The 'bench-mark' of the project is to reduce the roughness average of SL models to below 1.5J!m on all surface, without loss to the initial part geometry.

The 'bench-mark' of the project is to reduce the roughness average of SL models to below 1.5J!m on all surface, without loss to the initial part geometry. A number of diluent agents were also added to both epoxy XB5170 and acrylic XB5149 resins. A series of standard geometries were selected with vertical, horizontal and stepped faces, and the solution was applied to both cured and un-cured 'green state' components by dipping and painting. Applying resins to green state models has potential benefits as it could reduce the overall processing time of prototype manufacture, since UV curing need only be performed once following additive coating.Hexandiol-diacrylate was found to be the most successful diluent when applied to both acrylic and epoxy resins, as other diluents were seen to cause rapid coagulation. On exposure to UV post curing it was found that due to the separation of the monomer chains within the dilute solution, cross polYmerisation was not possible.

On exposure to UV post curing it was found that due to the separation of the monomer chains within the dilute solution, cross polYmerisation was not possible. The SL parts remained tacky irrespective of both curing temperature and time exposure to intense UV. The percentage mix of diluent was reduced from 50% by volume to 25% by volume, resulting in coagulation of the epoxy resin. Painting and dipping of parts was carried out with the acrylate solution however photopolymerization remained unsuccessful. Unfortunately due to the 'tackiness' of the coating following curing, it was not possible to measure coating thickness.A number of parts coated with the 25% thinned SL solution were also placed in a nitrogen enriched environment, and subjected to UV post curing. It was found that with the inhibition of oxygen during curing, full photopolYmerization occurred, resulting in a fully cured components.

It was found that with the inhibition of oxygen during curing, full photopolYmerization occurred, resulting in a fully cured components. The coating was capable of reducing surface roughness over a complex geometry although the viscosity of the coating resulted in blocked holes and features resulting in a loss of part geometry. Nitrogen curing with lower viscosity 50% diluent was not successful.

Nitrogen curing with lower viscosity 50% diluent was not successful. Further research is now being carried out to establish a suitable diluent resulting in full photopolymerization, yet maintaining a suitably low solution viscosity.EPOXY PRIMER COATINGSThe initial trials showed that resin coatings had potential for reducing the surface roughness provided that the coating thickness which is a function of the initial viscosity, can be defined and controlled, hence alternative coatings possessing good wetting and adhesion characteristics were investigated.Jaxacote 22.22 is a commercially available three part epoxy loaded primer coating usedmanufacture of glass fibre composites as a filling agent for the surface of damaged gel coats. coating contains both a filler and a low temperature exothermic agent, in addition to a thinning agent which can be added up to 75% by volume to reduce the coating viscosity.

coating contains both a filler and a low temperature exothermic agent, in addition to a thinning agent which can be added up to 75% by volume to reduce the coating viscosity. The coating is formulated to have excellent wetting and adhesion to epoxy substrates, resulting in a high sheen finish. It can alsoabraded following curing, to give a sub-micron surface roughness.99theA series of experiments were undertaken to apply both concentrated and thinned epoxy primer to acrylic and epoxy SL parts by painting, spraying and dipping, with the intention of reducing surface deviation yet retaining or improving part geometry (Figure 8). Initial findings showed that the coating in its concentrated state can reduce surface deviation by up to 70% when applied by dipping, but the coating viscosity led to significant loss to part geometry (Figure 9). Due to the high viscosity in the concentrated state spraying was not possible.

Due to the high viscosity in the concentrated state spraying was not possible. Geometric definition can be improved by brush application however surface roughness is accentuated by the effects of 'brush' markings and more importantly individually painting the parts will increase the overall processing time.Following investigation with the concentrated solution, a 50% thinned coating was applied to both acrylic and epoxy samples by dipping, painting and spraying. At the lower viscosity it was necessary to apply multiple coatings in order to reduce surface roughness by significant levels. On vertical and horizontal planes coatings applied by spraying were found to result in Ra values of 1.61lm and 2.21lm Ra, this being a reduction from 2.11lm and 5.61lm respectively. However, due to the thickness of the coating applied no significant difference was seen on stepped surfaces.

However, due to the thickness of the coating applied no significant difference was seen on stepped surfaces. At the thinned viscosity multiple layers ofthe epoxy primer applied by both dipping and painting were found to result in surface reduction of up to 50% on stepped planes. After two layers applied by dipping vertical and horizontal planes were reduced from 5.11lm and 2.11lm to 1.21lm and 0.61lm Ra respectively (Figure 10). The effect ofthe epoxy primer as seen in Figure 10 was not only a reduction in the Ra value of the component surfaces, but also improved geometry of the StereoLithography model nearer to the intended CAD profile.

These results show that thinned epoxy primer (Jaxacote) is capable of reducing surface roughness on parts of limited complexity with only minimal blocking of holes and features, although the time taken to build up sufficient coating thickness needs to be reduced.CONCLUSIONSIn conclusion it can be stated that, a number of positive directions for future research have been established. The use ofthinned SL resin as an additional coating, will if successful, result in near transparent SL models requiring no other post process finishing, yet retaining their intended geometry. The epoxy primer (Jaxacote) has been shown to be an excellent coating for SL components, either as a base for other coatings or as a method of surface finishing.

With the application of thinned primer by dipping it is possible to reduce surface deviation rapidly with minimal change to the part geometry.100ACKNOWLEDGEMENTSThe authors would like to thank the following bodies for their contribution to this research : EPSRC, Rolls Royce (Aerospace group), Britains Petite, Sonneborn & ReickREFERENCES1.Jacobs P.F "Rapid Prototyping & Manufacturing Fundamentals ofStereoLithography" Society of Manufacturing Engineers First Edition, pI 7, 1992Jacobs P F, Jouni P, Bedal B "Surface Quality" INSIGHT, 3D systems publication 1994.3.Jacobs P.F ''Quickcast™ 1.1 & Rapid Tooling" 4th European conference on Rapid Prototyping and Manufacturing, pp 1 - 27 Belgirate, ITALY, June 13th, 14th & 15th, 19954.Frank D "Preferred direction of build for rapid Prototyping processes" 5th International conference on Rapid Prototyping, pp 191 - 200 Dayton, USA, June 12th, 13th, 14th & 15th, 19945.Instituto Superior Tecnico "INSTANTCAM - Work Area 3 - Final report" BriteEuram Project No.

C "Surface finishing techniques for Rapid Prototyping" Society ofManufacturing Engineers - Technical paper PE93-169 Rapid Prototyping Conference Dearborn, Michigan, USA, May lIth, 12th & 13th, 19937.Boon Solid Free Form Manufacturing, 1st International user congress, pp 1 - 18 Technical University of Dresden, GERMANY, October 28th, 29th & 30th, 1993Chua C.K "Case studies of Jewellery Prototyping using the SLA"1011 -Horizontal Plane - 0.22 f.!m Ra 2 -Vertical Plane - 4.2 f.!m Ra 3 - 45° Plane - 24 f.!m RaAccurate Clear Built Epoxy Structure part (ACESTM), showing Vertical, Horizontal & 45° Stepped planes with O.25mm layers.Figure J4mmAcrylic Star Weave™ SL componentAcrylic Star Weave™ SL component following 4 hrs tumbling with ceramic media in a metallic finishing system.Figure 2102SL part built INSIDE CAD boundarySL part built OUTSIDE CAD boundaryADDITIVE COATINGABRASIVE FINISHINGCHEMICAL MACHININGThree phase approach to the post process surface finishing of SL componentsFifJUre 3Laser scan patterna = Layer thickness e = Plane angleParabolic curing profile of SL layer edgesFifJUre 4103Roughness Average -LA Ra = - - LWhere A is the area of profile above and below the line of the arithmetic mean, and L is the sample length taken along the line of the arithmetic mean.=An individual unit of area above or below the arithmetic mean line.

NITE ELEMENT ANALYSIS & STRAIN GAUGING OF THE STEREOLITHOGRAPHY / INVESTMENT CASTING SYSTEMRichard Hague and Phill Dickens Department of Manufacturing Engineering and Operations Management University ofNottingham Nottingham, UKABSTRACTMany metal parts have been produced from stereolithography (SL) models via the investment casting route. However, it is still not possible for every foundry to directly use SL models as thermally expendable patterns and gain the same success as achieved with wax patterns. Significant drawbacks still exist with the QuickCastTM structure that restricts its use to specialist investment casting foundries who are willing to alter their standard techniques.As part of a continuing work programme at the University of Nottingham, the stresses that are created in the SL/ceramic construction have been determined using simple stress analysis and finite element analysis techniques.

Further work has involved connecting strain gauges and thermocouples to SL parts in order to confirm the results obtained with the theoretical stress analysis. Inspection of the results obtained is aiding the generation of new build structures to enable the successful autoclaving of SL models.Details of the work to date are outlined in this paper, along with the results obtained.INTRODUCTIONSince 1993, with the introduction of the original QuickCastTM build structure 1, it has been possible to gain functional metal parts from stereolithography (SL) models using the investment casting process. The accuracy of the SL models lends itself well to investment casting because of the high tolerances that can be achieved when using this casting method. Thousands of metal parts have now been produced2 using the QuickCastTM process in a variety of different metals.However, there are significant drawbacks associated with using SL models as thermally expendable patterns.

The main problem area concerns the gross mismatch between the thermal expansions of the cured SL resin and the ceramic she1l3 . Because the epoxy resins used for the models are thermoset plastic materials, they do not melt under the influence of heat but continue to expand. It is this expansion that causes the catastrophic failure of the shells. The QuickCastTM build styles - effectively thin walled hollow parts with an internal open lattice structure of closed triangles (QuickCast1.0) or squares (QuickCast1.1) - were developed to allow the expanding SL parts some room to expand into and collapse during autoclaving.

Yet, experience has shown that when using conventional autoclaving techniques for dewaxing (still necessary because of the wax tree that the SL models are usually connected to), the QuickCastTM models still generally cause the shells to crack.523Due to these difficulties associated with achieving a casting from an SL model, it has generally been the more progressive foundries4 , more willing to adapt their existing techniques, who have been able to make the most of the stereolithography process for making one off or short-run metal components. However, it is desirable that all investment casting foundries should be able to take a stereolithography model and, by using their standard techniques (including the autoclave), be able to achieve a finished casting.Work at the University of Nottingham is continuing into the investigation of the stresses caused by the expansion of the SL models in the green ceramic shells as the construction is passed through the autoclave cycle.

This work is giving a much clearer understanding of what is happening to the SL / ceramic construction and is giving a clear indication of the necessary requirements for the design of 'successful' internal build structures.This paper is a continuation and update of the work presented previously5. It gives details of the stress analysis performed to date and indicates the future direction of the research programme.BACKGROUND WORKFailure of SL ModelsDue to the relatively large expansion of the cured epoxy resin, autoclaving solid SL models will cause the investment casting shells to crack.

To overcome this, QuickCastTM models were developed to enable investment casting shells to survive the autoclave phase by allowing the models to soften and then collapse in on their own voids 1 , Yet despite this, experience has shown that QuickCast™ models still generally cause the shells to fail in the autoclave phase6.Thus, to enable QuickCastTM parts to be used as thermally expendable patterns, it has been recommended7 that foundries skip the steam autoclave phase and dewax shells containing these models directly in the flash-furnace. Even though this method (coupled with other significant modifications to the standard investment casting process) works, it should not be viewed as a viable solution to the problems of ,dewaxing' QuickCast™ models.

What ideally required is a build structure that allows the direct substitution of the SL part for the wax, with out any variation to the standard investment casting process.On a brighter note, trials have shown5 that it is entirely possible, without modification, to autoclave thin walled hollow parts. This indicates that it is very likely that a structure(s) exists in between a hollow and solid that will also allow the autoclaving of parts (built in that structure) to be successful. If a stereolithography / investment casting construction can make it through the autoclaving stage there will normally be no problem in achieving the resultant casting.Sectioning of the autoclaved shells that contain these thin walled hollow parts shows that the SL models shatter within the confines of the shell.

On investigation, this appears to contradict the QuickCast™ 'theory' of the part softening and then collapsing in on itself.524If the parts were to soften and then collapse (as the 'theory' states), then upon sectioning of the autoclaved shells you would expect to see merely a buckling away from the walls of the SL model. The shattering of the hollow parts indicates that the stereolithography parts have indeed failed before they have time to soften. This transforms the reasoning behind the design of new structures, in that models must therefore be designed to fail (under the influence of their own expansion) before their expansion causes the ceramic shell to crack.Material Properties as a Function of TemperatureDetermining the changes in material properties as temperature increases, especially of the cured SL resin, is fundamental to understanding why investment casting shells containing SL models are failing.

The key to the failure of the shells lies in the high coefficient of thermal expansion (CTE) of the resin when compared to that of the green ceramic shell.The determination of the material properties is crucial to enable stress calculations to be performed. The relevant properties obtained are used in simple stress analysis tools (such as Lames theories8) and more complex finite element analysis (FEA) packages.A summary of the graphs for thermal expansion and Youngs modulus (Figures 1 & 2, respectively) are repeated in this paper for reference. Key points to note are the doubling in thermal expansion (from approximately 88 x 10.6 that occurs around the glass transition temperature of the material. Correspondingly, there is a significant drop in the Youngs modulus at temperatures above the glass transition.to 181 x 10.6)The graphs clearly show the sharp decrease in strength of the resin as the samples are ramped through an increase in temperature.

This indicates that most of the problems should be occurring at the lower temperatures, where the QuickCast™ models still have enough material strength not to collapse but enough expansion to crack the shells.STRESS CALCULATIONSFinite Element Analysis: Thick Walled CylinderTo complement the work previously performed using Lames theories to determine the stresses created in a simple stereolithography / investment casting construction, the Thick Walled Cylinder (TWC) problem was also modelled using an FEA package. (In summary, a TWC in this instance is taken to mean an open ended 'cylinder' of ceramic shell encasing a solid epoxy resin core. See ref. 5)Figures 3 & 4 show the results obtained using Lames theories and the FEA package, respectively for resin cores of 1mm and 50mm diameters. There is a clear correlation between the results obtained using these different theoretical methods.

There is a clear correlation between the results obtained using these different theoretical methods. The profile of the curves obtained from the FEA work match closely those obtained with Lames, except that there is a slight increase of temperature at which the maximum stress occurs with the FEA results. The magnitude of the stresses obtained with the two methods are also very similar to each other.525The similarities of the results from these two theoretical methods gives some degree ofconfidence as to suitability of using the FEA technique to model more complex examples.Finite Element Analysis: 50mm CUBEBasic theoretical stress analysis techniques, such as Lames theories, limit the complexity of the parts that can be analysed. Having confirmed the validity of the FEA package, the FE analysis was next extended to simulating the stresses induced in a SL 50mm cube encased in a shell 5mm thick; this represents a more complex geometry containing edges and comers.

Though it is envisaged that eventually cubes containing QuickCast™ structures will be modelled, the FEA was (for the purpose of this paper) restricted to cubes of varying hollows. Initially, three variation of the 50mm cube have been analysed; these include:50mm Cube; Solid • 50mm Cube; Hollow with a wall thickness of5.0mm • 50mm Cube; Hollow with a wall thickness of 2. 5mmA schematic of the FEA model is detailed in Figure 5. It is hoped that in future, thinnerwalls as well as 'structured' parts will be modelled.Figure 6 shows the stress profile over a range of temperatures for a solid 50mm block, a hollow 50mm block with a wall thickness of 5mm and a hollow 50mm block with a wall thickness of 2. 5mm. All of the resin blocks are covered with a shell 5mm thick.The introduction of comers have, predictably, had a great effect on the concentration and value of the stresses induced in the ceramic. For the solid 50mm block, there is a maximum stress of about 130MPa.

For the solid 50mm block, there is a maximum stress of about 130MPa. Comparing this to the value obtained from the 50mm diameter core in the thick walled cylinder example of about 16MPa, it is clear that the comer has had a dramatic effect.Though these hollow parts have been modelled with relatively thick walls, the effect on the induced stresses by introducing a hollow into a solid part is striking. There is a dramatic drop in stress as the wall thickness decreases. The block with a 2.5mm thick wall produces stresses in the ceramic approaching a value that is thought to be autoclavable. Interestingly, reducing the wall thickness has the same effect as reducing the resin thickness in the TWC problem, in that it decreases the temperature at which the maximum stress occurs (see Figure 4).

It should be noted that the maximum induced stress (in the ceramic) occurs at the mid-point of each edge of the cube.STRAIN GAUGINGInitial Strain GaugingOne of the principal reasons for performing the FEA was to identify not only themagnitude of the maximum stress, but also its location.526Initially a Solid 50mm stereolithography cube was built. This cube was then strain gauged and thermocoupled at the positions identified by the FEA work. An example of the strain gauging of the solid resin block can be seen in Figure 7. Having applied the gauges to the surface of the model, the initial face coat of the ceramic shell was applied and allowed to dry. With care, strain gauges were also applied to this initial face coat and the part was then coated with its subsequent ceramic layers, up to the normal thickness.This strain gauged block was then placed into an oven and the results were recorded. A sample of the results can be seen in Figure 8.

A sample of the results can be seen in Figure 8. This figure shows the values gained for one of the strain gauges attached to the resin block and the equivalent gauge attached to the ceramic shell. It seems that there is a clear break in the ceramic around 60°C. This initial result was very encouraging because it seems to confirm the observation that the majority of the problems are happening at or below the glass transition temperature of the resin. Figure 9 shows the shelled cube after it has gone through the heating process. It can just be seen that cracks have propagated from the centre of each edge which corresponds exactly with the predictions of the FEA.

It can just be seen that cracks have propagated from the centre of each edge which corresponds exactly with the predictions of the FEA. However, it is surprising that even though the ceramic shell has cracked under the influence of the heat, it has still remained mainly intact.One initial drawback that became apparent during this test is that to gain a result from thestrain gauged ceramic, it seems that a crack has to propagate directly over the strain gauge.These results appear to go some way to confirming the work performed with Lamestheory and the FEA work.Further Strain GaugingHaving achieved a reasonable result in the initial laboratory based 'dewaxing', the strain gauging was then expanded to encompass the full range of examples that were simulated using the FEA. For comparative analysis, three (3) of each type were fabricated. The types built were:Solid blocks • Hollow blocks, wall thickness 5.

For comparative analysis, three (3) of each type were fabricated. The types built were:Solid blocks • Hollow blocks, wall thickness 5. Omm • Hollow blocks, wall thickness 2.5mm(Referred to as Solid _1/2/ or 3) (Referred to as Hollow 5.0mm 1/2/ or 3) (Referred to as Hollow 2.5mm 1/2/ or 3)and also, to repeat previous work involving the autoclaving of thin-walled hollow partsHollow blocks, wall thickness 1.0mm(Referred to as Hollow 1. Omm_1/2/ or 3)As before, strain gauges and thermocouples were fixed onto the surface of the SL blocks and also attached to the initial 'face-coat' of the ceramic. For the Solid models an additional . thermocouple was also placed at the centre of the block.

For the Solid models an additional . thermocouple was also placed at the centre of the block. These blocks were then subjected to a real 'dewaxing' in a commercial steam autoclave.,A representative sample of the results obtained for the Solid block can be seen in Figure 10 and, similarly, an example of the results for a Hollow block (in this instance, Hollow5.0mm_l)527is shown in Figure 11. As expected, all the shells failed except for the ones containing the Hollow_1.0mm samples.DISCUSSIONThe clear result from the initial strain gauge work has not been duplicated in the autoclaving trials. There seems to be a stark difference in the results achieved when the autoclave is used to 'dewax' the samples. The graphs are far more difficult to interpret, but one of the most obvious thing to note is the conspicuous increase of temperature at which the shells appear to. fail.This apparent increase can be explained if the temperatures recorded in the Solid block are referred to.

fail.This apparent increase can be explained if the temperatures recorded in the Solid block are referred to. Figure 12 shows the temperatures recorded at the interface between the shell and the resin and also at the centre of Solid 2. The interface temperature has a very steep temperature gradient, but it can be clearly seen that the temperature at the centre of the block lags way behind this interface value. This demonstrates the insulative properties of the epoxy resin. Though the shells appear to be failing at about 120 to 130°C, it should be noted that this temperature is the surface temperature of the cube so, crudely speaking, the actual average temperature of the resin block at which the shells fail correlates to about 60 to 70°C. This is much more in line with the value predicted with the FEA package and demonstrated with the initial strain gauge trials.The heat input rate is clearly much more of an issue than was previously thought.

The autoclave that was used to 'dewax' these samples offered a very steep heating input, and consequently destroyed the shells (see Figure 13). The oven that was used to heat the solid block in the initial strain gauge work (described previously) provided a more gentle heating rate that allowed the resin cube to heat up more uniformly. The heating rate for the initial strain gauge work (again measured at the interface of the ceramic and model) is more in line with the rise in temperature recorded at the centre of the equivalent autoclaved block. (See Figure 14)In or.der to initially simplify the problem, the theoretical work performed (Lames theory and FEA) assumed that the SL / investment casting construction reached an instant and uniform temperature throughout. For the initial strain gauge work, because the heat input rate was relatively slow, the temperature distribution through the resin cube was more uniform.

For the initial strain gauge work, because the heat input rate was relatively slow, the temperature distribution through the resin cube was more uniform. It is for this reason that the initial strain gauge work reflects more closely the theoretical results. It should be noted that although the solid cube heated in the oven also failed, it did so far less catastrophically than the autoclaved version.For samples that were heated in the autoclave, the resin cube can be separated into two areas: an outside 'skin' that is getting progressively hotter and an inner core that, due to the temperature lag, initially remains mainly at ambient temperature. (The outer layer, of course, eventually consumes the inner as the effects of the lag are diminished). However, whilst the inner core remains solid, it restricts the softening and expanding outer layer from deforming.

However, whilst the inner core remains solid, it restricts the softening and expanding outer layer from deforming. So, as the outer 'skin' gets progressively hotter, its expansion increases to such a rate that it causes the shell to fail.528SummaryThe relatively novel approach to using strain gauges appears to present some difficulties. There are clear problems associated with using strain gauges that are imbedded and adhered to the ceramic with an adhesive that has a thermal expansion far greater than that of the ceramic itself. The problem of the expanding adhesive simply reflects the original problem of the cured SL models causing the shells to crack due to the gross mismatch of thermal expansion.SL parts that do not crack the shells during the autoclave phase (not published) show that the strain gauges are appearing to register a negative strain. This apparent compression of the ceramic runs entirely contrary to what actually happens.

This apparent compression of the ceramic runs entirely contrary to what actually happens. However, a logical explanation of this phenomena is that as the adhesive coating the strain gauge tries to expand under the influence of it is unable to because it is itself encapsulated by the ceramic shell. As the ceramic is heat, the strain gauge is forced into compression, and thus registers a (relatively) not expanding, negative strain.CONCLUSIONSThere are clear similarities between stress values obtained for the thick walled cylinderusing Lames equations and the results obtained with the FEA package.Though these results are preliminary and simplified, they clearly show a general trend, and support the idea that the majority of the problems are occurring before or about the beginning of the glass transition temperature of the cured resin.The initial strain gauge results that were performed in the low heat input oven offer good substantiation of the theoretical work performed.

The results obtained from the more aggressive autoclave trials are more confusing and seem to highlight the difficulties of using strain gauges in such an environment.ACKNOWLEDGEMENTSThe authors would like to thank 3D Systems Inc. and the UK Economic & SocialResearch Council (ESRC) for their help and financial support in this research programme.REFERENCESJACOBS, P.F. Stereolithography 1993: Epoxy Resins, Improved Accuracy & Investment 1 Casting. 2nd European Conference on Rapid Prototyping & Manufacturing, Nottingham, England, 15th-16th July, 1993. pp 95-113. Ed: P.M. DICKENS. ISBN 09519759 1 92 Prototyping & Manufacturing. 13th-15th June, Belgirate, Italy. 1995. pp 1-27. Ed: P.M. DICKENS. ISBN 0 9519759 4 3JACOBS, P.F. QuickCastTM 1.1 & Rapid Tooling. 4th European Conference on Rapid529WIMPENNY, D. Rapid Prototyping: An Overview of the Technology Applied to Investment3 Casting. 8th World Conference on Investment Casting. June 1993, p23-30.4 1995.

nite Element Analysis of Curl Development in the Selective Laser Sintering ProcessK.W. Dalgarno, T.R.C. Childs, I. Rowntree & L. Rothwell Department of Mechanical Engineering, University of Leeds, Leeds, LS2 9JT, UK.AbstractThe work reported within this paper is concerned with the development of analytical procedures which will allow the accuracy of parts generated by selective laser sintering to be predicted. One source of inaccuracy is curl, which results in curved part edges of flat plates manufactured lying horizontally in the part bed. This paper reports on the use of finite element techniques to model The models have been validated through comparison of f.e. results with the results of experimental builds, and extended to allow the influence of "bases" on the development of curl to be examined.the development of curl.1. IntroductionCurl in selectively laser sintered parts arises mainly from thermal distortion of parts within the build volume during processing.

IntroductionCurl in selectively laser sintered parts arises mainly from thermal distortion of parts within the build volume during processing. This results in nominally flat surfaces which lie horizontally in the part bed becoming warped. Figure 1 shows schematically how a nominally rectangular 90 x 26 rom, 50 layers thick polycarbonate plate made by SLS will exhibit curl.Figure 1 - Effect of Curl on Part GeometryPrevious work [1] has identified different regions of curl within a part. A microscopic amount of curl exists as a result of the definition which can be expected from the process. Thereafter the curl can be divided into two effects. The first of these appears over approximately the 2.5 rom of part closest to the edge of the part and on parts of with the geometry shown in Figure 1 has a radius (R1 in Figure 1) of around 20 mm.

The first of these appears over approximately the 2.5 rom of part closest to the edge of the part and on parts of with the geometry shown in Figure 1 has a radius (R1 in Figure 1) of around 20 mm. The second extends from the end of R1 approximately 10 rom further towards the centre of the part and has a radius (Rn in Figure 1) of around 60 rom. R1 curl has been seen to exist from the sintering of the first layer in part build, while Rn curl seems to develop as layers are added to the part. This paper reports on the use of the finite element method to analyse the development of Rn curl in rectangular polycarbonate plates made by the SLS process on a DTM Sinterstation 2000. One particular area of interest was the localised nature of the curl: previous work had failed to identify any reason why the curl should be confined to the part edges.559Previous work also investigated the effect of bases on Rn curl.

Figure 2 shows results from an experimental investigation into the effect of the power input to the base, [P/(US)]base, on the resulting Rn curl. The graph in Figure 2 suggests that bases have no effect on Rn curl until the base is sufficiently dense to offer some restraint, but that thereafter no additional benefit is derived by further increasing the density of the base. The reasons for this were not fully understood and a further aim of the finite element work was an investigation of this effect.200Figure 2 - Effect of Energy Input to Base on RII Curl. From [1].O'------L.--'---.l.-.---I.-....I----I 0.060.020.04[P/(Us)] base' J / mm 2The finite element method has previously been used by Bugeda et al [2] to study curl development in the stereolithography process. They concluded that curl in stereolithography arose from volumetric shrinkage of the resin when cured. There are seen to be two mechanisms for curl development within selectively laser sintered parts.

There are seen to be two mechanisms for curl development within selectively laser sintered parts. The first of these is shrinkage as a result of the sintering process, and the second is a thermal strain which results from powder to form the next layer being deposited on the most recently scanned layer. The fresh powder layer is at a lower temperature than the scanned layer, and so will produce some thermal contraction within the scanned layer. Both of these mechanisms result in compressive strains in the layers as they are built. The approach taken within the work reported here has been to use the finite element method to model part build by adding layers of elements to represent the addition of material to a part, and introducing compressive strains into the layers as they are added.2.

Finite Element Model of Part BuildThe finite element solver used within the work described here was ABAQUSIStandard [3].2.1 GeometryFigure 3 shows the finite element mesh used within the analysis. The mesh represents the first 10 layers of the 90 x 26 mm rectangular polycarbonate plate part described above. All layers excepting the first two have a thickness of 0.125 mm, which was the default layer thickness within the manufacture of the parts. The first layer has a thickness of 0.66 mm, and the second a thickness of 0.25 mm. These thicknesses are based on measurement of layer thickness on one and two layer parts with the same cross sectional area. Further measurements indicated that subsequent layers were the expected 0.125 mm deep. The additional thickness in these first two layers is as a result of "bonus z", whereby additional material is sintered early in the build as a result of large thermal penetration into the part bed.

The additional thickness in these first two layers is as a result of "bonus z", whereby additional material is sintered early in the build as a result of large thermal penetration into the part bed. The model was symmetric about the centre of the plate, and so only half the plate has been modelled. 2D analysis was used as it was the distortion along the 90 mm length of the plate that was of interest.560Il.-_2L1:t=tt= !r--Layer10Figure 3 - Finite Element Mesh--""- Laye -....t:> Laye .... "'.r--Layer3 r2r 1Rigid SurtaceThe elements used within the analysis were four noded coupled temperature-displacement plane stress elements. Coupled elements allowed for the compressive strain to be introduced to the layers as they were built by artificially lowering the temperature locally within the mesh. The procedure used was to initially define the elements for all ten layers that would be built.

The procedure used was to initially define the elements for all ten layers that would be built. The elements for layers two to ten were then removed from the analysis (using the *MODEL CHANGE card within ABAQUS) and re-introduced layer by layer as the analysis progressed.In practice the surface is unlikely to be rigid, but if the part bed is assumed to remain at the same temperature throughout the sintering process it is likely that the bed will be stiff relative to the part.The part bed has been modelled as a rigid surface.2.2 Material PropertiesThe parameters used in building the parts produced parts with a density of 700 kg/m3 . The modulus of the polycarbonate within the analysis was assumed to be 50 MPa, with a Poisson's ratio of 0.4. There is no real scientific basis for the use of an elastic material model, or for its value.

There is no real scientific basis for the use of an elastic material model, or for its value. In the absence of better material property data a "smeared" elastic response has been assumed in order to evaluate the usefulness of the finite element method in modelling the selective laser sintering process. Further work is planned to gain more information on material property values through the sintering process.The interaction between the part and the part bed was modelled with a coefficient of friction of 1. This high coefficient of friction value was chosen to reflect the fact that the real surfaces would be quite rough, giving high resistance to relative movement.2.3 Boundary Conditions and LoadingThe first step within the analysis was to analyse the first layer. The only boundary condition applied was to enforce symmetry; the nodes at the centre of the part were therefore constrained not to move in the 1 direction (shown in Figure 3).

The only boundary condition applied was to enforce symmetry; the nodes at the centre of the part were therefore constrained not to move in the 1 direction (shown in Figure 3). A gravity load was applied as a body force to561all elements active in the model. The compressive strain was then imposed on the layer by artificially reducing temperature within the model. The strain imposed using this method was 2.6 x 10-4. Initially larger values for the developed strain were used, on the basis of calculations by Childs et al reported in [l], but these values were found to overestimate the amount of curl. It may be that initially large strains in the layers may be somewhat relaxed by temperatures developed in a layer while subsequent layers are being scanned.

It may be that initially large strains in the layers may be somewhat relaxed by temperatures developed in a layer while subsequent layers are being scanned. After the first layer has been loaded in this way the elements making up the second layer were introduced, and the gravity load and compressive strain applied to them, with this procedure repeated until all ten layers of the part had been introduced and loaded. Figure 4 shows a detail of the resulting displaced mesh.Figure 4 - Displaced Mesh2.4 ResultsIn order to validate the analysis the results have been compared with the measurements of experimental builds. Table 1 shows how the results from the finite element model compare with measurements of distortion on two and ten layer parts of the same geometry. The data points used are away from the outer edge of the part to avoid confusion with R1 curl.

The data points used are away from the outer edge of the part to avoid confusion with R1 curl. The results shown in Table 1 were taken as indicating that the distortions predicted by the finite element model were of a similar order to those exhibited by the parts.The results overall suggested that the combination of gravity loads, the strain induced within layers as they are manufactured, and part/bed interaction leads to Rn curl being seen only close to part edges.Measured displacement in 2 direction 3.25 mm from part edge on two layer part FE predicted displacement in 2 direction 3.25 mm from part edge after two layers built Measured displacement in 2 direction 5.5 mm from part edge on ten layer part FE predicted displacement in 2 direction 5.5 mm from part edge after ten layers built0.025 mm0.011 mm0.05 mm0.052 mmTable 1 - Comparison Between FE Predictions and Measurement of Distortion5623.

Analysis of the Effect of BasesFigure 2 shows that the effect of bases on Rn curl is dependent on the power input per unit area to the base. The power input effectively defines the density developed within volume being scanned by the laser. It is assumed that the region in Figure 2 where the base has no effect is due to the lightly sintered powder being unable to sustain any load (and therefore acting in the same In order to investigate the results shown in Figure 2 analyses with manner as the powder bed). bases appropriate to [P/(US)]base values of 0.035 and 0.06 J/mm2 have been carried out, to try to identify why no further improvement occurs beyond that shown for a [P/(US)]base value of 0.35 J/mm2.

bases appropriate to [P/(US)]base values of 0.035 and 0.06 J/mm2 have been carried out, to try to identify why no further improvement occurs beyond that shown for a [P/(US)]base value of 0.35 J/mm2 . These two values of [P/(US)]base would be expected to generate base densities of 580 and 700 kg/m3 respectively (the higher value being a base with the same density as the part being built).3.1 Model DevelopmentThe mesh used for analysing the effect of a base density of 580 kg/m3 is shown in Figure 5. The model operated in essentially the same way as that described for part build. The base used in the analysis is four layers thick, corresponding to those used to carry out the measurements shown in Figure 2. The depth of the first two layers in the build (in this case the first two layers of the base) has been changed as the "bonus z" effect increases with sintering power.

The depth of the first two layers in the build (in this case the first two layers of the base) has been changed as the "bonus z" effect increases with sintering power. Measurements on 50 layer parts gave "bonus z" values for a part made with a laser power of 0.035 J/mm2 as around half of that for a part made with a laser power of 0.06 J/mm2 . The depth of the first two layers was therefore halved. The layer thickness in the model for a base with a density 700 kg/m3 was the same as for the analysis described in section 2.Layer 10Layer 2I !~, V ,/Layer 1- - - ......> Base 1 ."",..Base 4Figure 5 - Finite Element Mesh of Part Build with Base. Base Density 580 kg/m3563.In both cases the base was extended 2 nun beyond the part in the 1 direction. This is the radius of the base (which is cylindrical) would have been. The approximately what cylindrical shape of the base means that a 3D analysis would have been necessary to reflect the part geometry completely.

The approximately what cylindrical shape of the base means that a 3D analysis would have been necessary to reflect the part geometry completely. A 2D analysis was retained for computational efficiency, although this means that the effect of a large area of base material is lost. The effect that the area of the base has on part curl was investigated by carrying out one further analysis with a base which extended 7 mm beyond the part. This analysis was carried out with a geometry and material properties appropriate to a base with a density of 700 kg/m3After the base and two layers of part contained within the base have been completed the addition of subsequent layers deposits powder on top of the protruding part of the base. The effect this has was modelled by applying a pressure load to the top of the elements of the fourth base layer.When bases are used the layers which occupy space within the base cylinder and the part volume are scanned twice by the laser.

The laser first rasters the base layer, with the laser power per unit area appropriate for the base, before rastering the part layer, again with the appropriate It has been assumed within the analyses that the second raster causes laser power per unit area. the strain developed during the first raster to be relieved (initial analyses suggested that if this were not the case the part made with a 700 kg/m3 density base would curl more than the part made with no base at all). For the 700 kg/m3 density base it has been assumed that all of the developed strain is relieved by the second raster. For the 580 kg/m3 base it has been assumed that amount of strain developed during the first raster is proportional to the amount of densification which takes place.

For the 580 kg/m3 base it has been assumed that amount of strain developed during the first raster is proportional to the amount of densification which takes place. When the second raster occurs this strain is assumed to be relieved, but a further strain imposed, as the density of the layer goes from 580 kg/m3 to 700 • By this reasoning a strain of 1.46 x 10-4 is assumed to be developed within layers 1 & 2 kg/m3 in the model of the 580 kg/m3 base build. The modulus of the 580 kg/m3 base within the model was 16.7 MPa, experimental measurement of the room temperature elastic modulus of polycarbonate parts made by laser sintering suggested that a 580 kg/m3 part would have one third of the modulus ofa 700 kg/m3 part. The modulus value for the 580 kg/m3 dense material was therefore taken as one third of that used for the 700 kg/m3 dense material. The Poisson's ratio was assumed to be 0.4.

The modulus value for the 580 kg/m3 dense material was therefore taken as one third of that used for the 700 kg/m3 dense material. The Poisson's ratio was assumed to be 0.4. To further investigate the effect of the second raster releasing the strain in the layer as sintered one further analysis was carried out. This analysis used the same mesh as for the part built without a base, but assumed that the first two layers were rastered twice.

This analysis used the same mesh as for the part built without a base, but assumed that the first two layers were rastered twice. The first two layers were therefore assumed to have no developed strain.Other than the points mentioned above the analysis procedure, material properties andloading were the same as that for the part without a base described in section 2.3.2 Results and DiscussionThe results from the four analyses can be seen in Figure 6, where the position of the lower surface of the part after two and ten layers have been built within each of the analyses is plotted together with the results obtained for the analysis of a part without a base.5640.025No base-700 kg/mA347 mm base---E E---c0.02700 kg/mA3 53 mm base580 kg/mA3 47 mm base0 +:i u Q) l.....

The results shown for the analysis of a 700 kg/m3 density base and a 580 kg/m3 density base do not565(a)45(b)I45differ significantly from those of the part built without a base. The results for the analysis carried out where an additional 5 mm of base material were added show a significant reduction in the amount of curl generated. This would suggest that the additional area of cylindrical base not taken into account by a 2D model would have an effect on the amount of curl generated.(ii) The analyses with the 580 kg/m3 and 700 kg/m3 density bases show very little difference in terms of the developed Rrr curl. The residual strain in the first two layers of the part in the case of the 580 kg/m3 density base is greater than that for the 700 kg/m3 base, however this seems to be offset by this strain being applied to material with a lower density, modulus and (because of the effect of "bonus z") depth.

This provides some insight into the effect shown in Figure 2.(iii) The results suggest that sintering the first two layers twice to relieve the strain developed in these layers would reduce the amount of curl developed. As the results of the analyses of parts with bases is thought to underestimate the effect that the base has, further work will be required to assess whether or not the benefit derived is as great as that from using a base.4. ConclusionThe finite element method has been shown to be of value in understanding curl in selectively laser sintered components. The characteristic shape of components exhibiting curl is considered to arise from the interaction of strain developed in the layers of the part during the build process and gravity forces. The effect of bases on Rrr curl has also been analysed, and qualitative agreement between analysis and experiment shown.

The effect of bases on Rrr curl has also been analysed, and qualitative agreement between analysis and experiment shown. The work has made major assumptions with regard to material properties and viscosity, and further work will address these issues.References1. Berzins, M. et aI, "Densification and distortion in selective laser sintering of polycarbonate parts", in Solid Freeform Fabrication Symposium 1995, pp 196-203, ISSN 1053-2153, Proc. of the Symp. held in Austin, Texas on Aug. 7-9,1995, Pub. The Univ. of Texas at Austin.2. Bugeda, G. et aI, "Numerical analysis of stereolithography processes using the finite element method", Rapid Prototyping Journal, pp 13 - 23, Vol. 1, No.2, 1995.3. ABAQUS/Standard Users Manual, Version 5.4, Hibbitt, Karlsson & Sorenson Inc.566.

nite Element Analysis of Residual Stresses andDeformations in Direct Metal SLS ProcessWei Jiang School of Mechanical Engineering, Dalian University of Technology, 116024, P.R. ChinaK.W. Dalgarno and T.H.C. Childs School of Mechanical Engineering, University of Leeds, Leeds, LS2 9JT, U.K.AbstractDirect metal SLS processes use a high power laser beam to selectively fuse fully metallic powder to directly produce functional metal components or tooling. As a thermal dynamic process, it inevitably causes thermal stresses, stress-induced deformations and cracking. Understanding these is very important in controlling residual stresses and stress-induced deformations.In this research, the residual stresses and deformations in direct laser sintering of stainless steel is investigated by an integrated thermal and mechanical model. Temperature-dependent material properties are taken into account in both models.

Temperature-dependent material properties are taken into account in both models. Using a commercial finite element software, with sintered geometry and temperatures imported from a thermal model, the residual stresses and deformations of direct SLS of stainless steel are predicted. In the long term this will make it possible to achieve minimum residual stresses and deformations by controlling the SLS process parameters, material properties and other relevant parameters.1. INTRODUCTIONDirect metal SLS processes use a high power laser beam to selectively fuse fully metallic powder to directly produce functional metal components or tooling [1,2]. It has the capability of reducing cycle time in product development, lowering cost, and building a very complex geometries. However, as a thermal dynamic process, SLS process inevitably causes thermal stresses, stress-induced deformations and cracking.

However, as a thermal dynamic process, SLS process inevitably causes thermal stresses, stress-induced deformations and cracking. Understanding these is very important in controlling residual stresses and stress-induced deformations and in developing SLS into industrial grade rapid manufacturing method.Direct metal SLS processes are very complicated. During the process, metal powders under the laser beam undergo a phase change from solid to liquid and then back to solid within a very short time. This phase change is associated with latent heat, which has a significant effect on the temperature and residual stress distribution. Also the elevated temperature gradients present in this area both during the heating and the cooling, along with the sharp decrease in mechanical340properties during heating, yield non-homogeneous permanent strains and residual stresses after processing.

Residual stress has a great influence on structural strength, loss of edge tolerance and inter-layer debonding.Modeling direct metal SLS is a very challenging task. Although some SLS modelling work have been reported [3-7], it has concentrated on amorphous polymers. A thorough survey of the existing literature indicates that scant attention has been paid to the numerical modeling of direct metal laser sintering process. This paper reports finite element modeling of direct metal SLS. The residual stresses and deformations in direct metal laser sintering are investigated by an integrated thermal and mechanical model. In the thermal model, the nonlinear heat conduction equation with a moving Gaussian heat source is applied, and latent heat due to phase change is considered. The temperature-dependent material properties are taken into account in both models.

The temperature-dependent material properties are taken into account in both models. By using ABAQUS commercial finite element software, with fully sintered geometries and temperatures imported from thermal model, the residual stresses and deformations in the direct SLS of 316 stainless steel are predicted. The objective of this work is to understand residual stresses build-up and to develop a process simulation to optimize the SLS process and to control it intelligently.2. MODELLING 2.1 MethodThe strategy to solve problems of thermally induced residual stress is to link heat transfer analysis to stress analysis, by an integrated thermal-mechanical model. Fig. 1 presents the analysis procedures.In the thermal model, the governing principle of SLS process is modeled as non-linear heat conduction equation with a moving Gaussian heat source [8]. During direct metal SLS process, the powders undergo a phase change from solid to liquid and then back to solid.

During direct metal SLS process, the powders undergo a phase change from solid to liquid and then back to solid. This phase change is associated with latent heat, and the temperature recovery method [9] is employed to take account of this. Temperature dependent thermal conductivity and specific heat are considered in the thermal model. In order to predict the density field and sintered geometry, the resulting temperature/time history is linked to a liquid-phase sintering law and mass conservation principle [10]. The shrinkage due to solidification is assumed only in the z direction, as shown in Fig. 2(a).The transient temperatures, sintered geometry and density field are output from the thermal model. After filtering out unnecessary information by interface programs, the temperatures below the melting point and geometries of the fully sintered region are used as input to the mechanical model to predict the evolution of residual stresses and deformations.

The thermal meshes are discretized into four-node brick elements, as shown in Fig. 2(a). Each four-node brick is regarded as a pair of three-node triangles. In order to accurately predict the large gradients of temperature near the heat source, the thermal mesh is relatively fine in the neighborhood of laser beam qlaser, its y-spacing is 1/6 of the beam diameter d (0.183 mm), and its z-spacing is 1/6 of the powder layer thickness tlayer, (0.083 mm). The mesh coarsens with distance from the heat source in multiples of this fine mesh size. The powder bed mesh is of the same size as the fine thermal mesh. Linear interpolations are used to map between the powder bed meshes and342the thermal meshes. The thermal meshes superimpose on the powder bed, moving at v in the Y direction by d/6 in each time step.Since partially sintered regions will not support thermal stresses, they were filtered out by an interface program.

The geometries of the fully sintered region whose density is greater than 7800 kg/m3 were imported to the mechanical model. The meshes in the mechanical model are of the same size as in the thermal model, that is 1/6 of laser beam diameter d in x-spacing and 1/6 of the powder layer thickness tlayer in y-spacing, as shown in Fig. 2(b). The fully sintered part is modelled using four noded plane stress element (CPS4R). The part bed is modelled as a rigid surface. With a reference node, the rigid surface is modeled as planar rigid surface elements, by IRS21 element.

3(a) [11]. In contrast to the heat conduction problem of a solid materials, the thermal conductivity, k, is also a function of temperature and local bed porosity ε, as shown in Fig. 3(b) [9]. The material is modeled as having liquidus temperature of 1380(cid:176)C, a solidus temperature of 1280(cid:176)C, and a latent heat of fusion of 308 kJ/kg. The heat transfer coefficient, h, was 5.3 W/m2K. The powder bed temperature, Tb, is 20(cid:176)C.The adopted idealization of the behavior of stainless steel under direct metal SLS process is assumed elastic for temperatures below melting point.

2(a).In mechanical model, the part bed in the analysis is assumed as a rigid surface. The interaction between the part bed and the powder was modeled with a coefficient of friction 0. Although in practice the surface is unlikely to be rigid, it is still assumed to be relatively stiff to the part, and to remain at the same temperature through out the analysis.A gravity load is applied to all the elements active in the model as a body force. The stress analysis model use the temperature profiles obtained in the heat transfer analysis to define the3441thermal loading. Since any change of temperature above the melting point will not produce any stresses, all the temperature beyond the melting point are filtered out by another interface program before it is imported into the mechanical model.3. RESULTS AND DISCUSSION 3.1 Temperature ProfileFig. 4(a) shows the transient temperature history as the laser beam moves.

4(b) shows the thermal stress history within the 0.0833 mm depth of material on the top of the sintered part. For σxx at all points the part first experiences a slight compression stress, then, as the beam arrives at this point, the metal melts. After the laser beam has past by, cooling occurs, and tensile stresses are generated. The tensile stress becomes greater as the melted metal(cid:146)s strength recovers. Fig.

The first is shrinkage as the result of the sintering process, the second is the result of the thermal stress-induced deflection. These deformation will cause parts warping (curvature) and loss of edge tolerance. Fig. 5 shows the development of these deformations.The sintered geometry is the coordinates imported from thermal model after filtering out partially sintered region. Thermal loading in the mechanical model will cause deflections in x and y directions. The total deformation in vertical direction is the sum of shrinkage due to solidification and y deflection due to thermal loading. Compared with thermal y deflection, the sintering shrinkage is the main reason for vertical warping and edge tolerance. As the heat source moves along x direction, the thermal deflections in y direction do not change significantly, while in x direction, the thermal deflections increase. After cooling down i.e.

As the heat source moves along x direction, the thermal deflections in y direction do not change significantly, while in x direction, the thermal deflections increase. After cooling down i.e. at time 2.562s, the x deflection becomes smaller.4 SUMMARY AND CONCLUSIONSThis research employs finite element method to evaluate the residual stresses anddeformations in direct metal SLS process. Three elements of analysis have been involved:Modeling the heat transfer problem using a FORTRAN finite element program, withconsideration of latent heat and temperature dependent material properties.Using interface programs to obtain temperatures below the melting point and the geometry ofthe fully sintered region, then using them as input to the mechanical modelModeling stresses and deformations as a function of time and location for direct metal SLSprocess using a commercial f.e.

code.From the simulation results we can conclude:Excluding edge effects, every position experiences a similar temperature history. • The deformation in the vertical direction is mainly caused by shrinkage due to the sintering of porous powders, while the deformation in horizontal direction is mainly the result of thermal loading.ACKNOWLEDGEMENTWEI JIANG gratefully acknowledge financial support from the China Scholarship Council under Grants 99821129 and National Science Foundation of China under Grants 59935110.REFERENCES1. JJ Beaman, JW Barlow, DL Bourell, RH Crawford, HL Marcus, KP McAlea. Solid freeform fabrication: a new direction in manufacturing; Dordrecht London: Kluwer Academic347Publishers, 1997.2. PF Jacobs, Stereolithography and other RP&M technologies: from rapid prototyping to rapid tooling: Society of Manufacturing Engineers in cooperation with the Rapid Prototyping Association of SME New York: ASME Press, 1996.3.

nite Element Analysis of the SALDVI ProcessK. Dai, J. Crocker, L. Shaw and H. MarcusDepartment of Metallurgy and Materials Engineering Institute of Materials Science University of Connecticut, Storrs, CT 06269AbstractSelective Area Laser Deposition Vapor Infiltration (SALDVI) is a developing solid freeform fabrication (SFF) technique aimed at the direct fabrication of ceramic and ceramic/metal structures and composites. SALDVI uses a layer-by-layer approach in which layers of powder are densified with solid material deposited from gas precursors by chemical vapor deposition (CVD) using laser heating. In this work, we have performed numerical simulation using the ANSYS code with 3-dimensional coupled field elements to calculate the temperature field and the part geometry resulting from the SALDVI process. The effects of the powder and vapor deposited material properties on the temperature distribution and the part geometry have been investigated.

The effects of the powder and vapor deposited material properties on the temperature distribution and the part geometry have been investigated. The result from the numerical simulation is found to be consistent with those obtained from experiments performed using the silicon carbide forming gas precursor Si(CH3)4 and SiC powder particles.Keywords: Solid freeform fabrication, Finite element modeling, Thermal analysis, SALDVI, Laser densification.IntroductionRecent advances in physical prototyping allow the production of freeform solid objects directly from a computer model without part-specific tooling and human intervention. These technologies have been termed Solid Freeform Fabrication (SFF). The basic feature of all the SFF techniques is their capabilities to build parts point-by-point and layer-by-layer selectively so that complex-geometry components and systems with highly integrated multifunctions can be fabricated in a single operation.

This new manufacturing technology can be used to the direct fabrication of ceramic and ceramic/metal structures and composites. To produce fully functional structural components, powder based approaches to SFF seem to be the most promising [1].Selective Area Laser Deposition Vapor Infiltration (SALDVI) is such a developing powder-based SFF process in which porous layers of powder are densified by infiltrating the pore spaces with solid material deposited from a gas precursor during laser heating [2]. SALDVI experiments have shown that the size and geometry of the densified part depend in large part on the temperature distribution in the SALDVI workpiece among other parameters [3]. The understanding of the temperature distribution in the SALDVI process is very important to achieve the insight into the effect of various process parameters on the shape of the SALDVI workpiece.

The understanding of the temperature distribution in the SALDVI process is very important to achieve the insight into the effect of various process parameters on the shape of the SALDVI workpiece. During the SALDVI process the relative density of the workpiece continuously changes with processing time until it reaches near full density. Because of such continuous changes in the relative density and thus the continuous change in thermal conductivity of the SALDVI workpiece, the transient temperature field of the SALDVI workpiece is too complex to393calculate by analytical methods. Therefore, finite element method (FEM) is used in this paper to quantify the transient temperature distribution and develop fundamental understanding of the formation of the dense part during the infiltration and densification process of the powder bed.

Such finite element analyses will certainly provide guidelines for controlling the SALDVI process to fabricate the part with desired geometry and dimensions. Furthermore, these analyses will also be useful for understanding of other SFF processes involving powder-to-dense-body transition such as selective laser sintering (SLS) [4].Finite Element ModelThe goal of this modeling effort is to simulate the densification process of powder bed and evaluate the effect of laser processing conditions on the size and geometry of the dense body resulting from the densification process. The finite element model used for this purpose is shown in Figure 1 and calculated using the ANSYS 3-dimensional coupled-field element (Solid5). The powder bed before densification is assumed to have a dimension of 14-mm length, 7-mm width and 5-mm height.

The powder bed before densification is assumed to have a dimension of 14-mm length, 7-mm width and 5-mm height. Thus, the powder bed in this study is very thick in comparison with the layer- by-layer fabrication technique, which typically has a powder layer thickness less than 0.5 mm. However, the present simulation allows for assessment of building the first solid layer from a thick powder bed rather than from a thin powder layer on the top of a substrate of different material. Densification of thin powder layer will be modeled in the future effort.Starting PointIFinishing PointYIZXPowder BedCross Section I- IFigure 1. Finite element model of the SALDVI process.As shown in Figure 1, the SALDVI process is modeled by moving the laser beam from one end of the powder bed to the other and monitoring the temperature change in the powder bed.

To simulate the Gaussian-shaped 1-mm diameter laser beam and the infiltration temperature used in the SALDVI experiment, the temperature of the laser beam is fixed at 1353K and the size of the laser beam is 1· 1 mm. Since the size of elements in the powder bed at the X-Z plane is3940.5· 0.25 mm, there are eight surface elements exposing directly to the laser beam at any given time during the infiltration and densification process. The scanning rate of the laser beam is assumed to be 2.5m m/s and thus the laser beam moves one element distance (0.5 mm) every 200 seconds in the X-direction. The powder bed before laser densification is assumed to have 50% of its theoretic density. When the temperature of an element in the powder bed reaches 1200K, the density of that element is changed to 80% of its theoretic value. It becomes 100% dense when the temperature of the element reaches 1273K.

It becomes 100% dense when the temperature of the element reaches 1273K. The part being densified is assumed to be in contact with a gas and the heat loss through the gas is approximated through the inclusion of a convection heat transfer coefficient (6· 10-5 w/mm2K) between the part and the gas. The temperature of the gas is assumed to be 300K and no temperature dependence of the convection heat transfer coefficient is considered.The properties of the material used in this simulation are summarized in Table 1 [5] and Table 2 [7-8].

Air (curve 19) [6] **dense cubic SiC (curve 3) [7] ***dense a SiC (recommended values) [8]Results of Finite Element SimulationThe temperature distribution on the surface of the part after one pass of laser scanning from one end to the other is shown in Figure 2. As expected, the temperature of the material within the laser scanning path is higher than that of the rest of the material. Furthermore, the temperature gradient in the material within the scanning path is smaller than that in the material outside the scanning path. This is due to densification of the material within the scanning path,395thereby the increased thermal conductivity and low temperature gradient within the scanning path.The shape of the part that is 100% dense resulting from the same scanning condition as that shown in Figure 2 is presented in Figure 3. It is noted that the cross section of the densified part can be divided into three regions. Region A represents the initial transient region.

It is noted that the cross section of the densified part can be divided into three regions. Region A represents the initial transient region. Region B is the steady state region for the processing condition and the material used in the modeling. Region C is the final transient region. The thickness of the part varies from region to region. Such phenomenon is related to either the change in the thermal conductivity when the powder compact changes to the dense solid, or the difference in heating time. Recall that under the current the elements at the two very ends are only subject to 200 seconds of laser heating, whereas all the other elements between the two very ends have exposed to the laser heating for 400 seconds. Thus, the thinner sections in Figure3 at the two very ends compared with their neighboring areas are due to shorter heating time.YZXFigure 2.

Thus, the thinner sections in Figure3 at the two very ends compared with their neighboring areas are due to shorter heating time.YZXFigure 2. Temperature distribution of the surface of the part after one pass of laser scanning from one end to the otherlaser scanning conditionsA B CSide ViewTop ViewFigure 3. The shape of densified part after one pass of laser scanningThe thickness change between the two very ends can be rationalized as follows. At the beginning of the laser scanning, the laser-heated area is surrounded by powder and heat dissipation is slow. Thus, more heat is available for densifying the powder under the laser beam, thereby a thicker densified section at the beginning of the scanning. As the scanning proceeds, the part size increases and more heat is dissipated through thermal conduction along the previously densified section. Thus, the thickness of the currently densified section becomes thinner than that of the previously densified section (e.g.

Thus, the thickness of the currently densified section becomes thinner than that of the previously densified section (e.g. the last part of Region A vs the thicker part of Region A). Finally, a steady state is reached that heat conduction through the previously densified section does not change much with time (Region B). Since more heat is conducted away through the previously densified section in the steady state region than that in the initial396transient region, the thickness of the section in the steady state region is smaller than that in the initial transient region.The temperature distribution along the cross section I-I of Figure 1 as a function of the laser scanning position is shown in Figure 4 which confirms the aforementioned reasoning. It can be seen that the region with high temperatures (i.e. > 1240K) is thicker during the initial scanning period except the first scanning point, and becomes thinner after that.

It can be seen that the region with high temperatures (i.e. > 1240K) is thicker during the initial scanning period except the first scanning point, and becomes thinner after that. The shape of the part obtained by the simulation shown in Figure 3 is similar to that obtained by the experiment [9].This study indicates that in order to fabricate a part with minimum transient regions the infiltration and densification temperature during laser scanning should be adjusted according to heat dissipation as well as energy input. This finding is different from the conventional wisdom that the laser input power should be adjusted to maintain a constant densification temperature during laser densification. This study shows that constant densification temperature produces a constant cross section only when a steady state of heat dissipation and energy input has been reached.

This study shows that constant densification temperature produces a constant cross section only when a steady state of heat dissipation and energy input has been reached. At the transient regions the infiltration and densification temperature should be adjusted if the transient regions are to be minimized.(a) (b)(c) (d) Figure 4. Temperature distribution in the cross section I-I shown in Figure 1 with (a) at the first scanning point, (b) at about 1/3 scanning distance, (c) at about 2/3 scanning distance, and (d) at the end scanning point397Conclusions1. An initial and a final transient region are present in the SALDVI workpiece. The thickness of the steady state section is smaller than that in the initial transient section. These phenomena arise from a combined effect of a constant infiltration and densification temperature and a heat dissipation rate that varies with the size of the SALDVI workpiece.2. Constant surface temperature (e.g.

Constant surface temperature (e.g. infiltration and/or densification temperature for powder bed) during laser scanning does not provide constant cross section. At the initial and final transient regions not only the laser power but also the densification temperature should be adjusted to minimize the transient regions. Acknowledgement – The authors gratefully acknowledge financial supports provided by the National Science Foundation under Grant No: DMI-9908249 and the Office of Naval Research under Grant No: N00014-95-1-0978.References[1] Ashok V. Kumar, Powder Deposition and Sintering for a Two-Powder Approach to Solid Freeform Fabrication, Proceedings of Solid Freeform Fabrication Symposium, David L. Bourell et al. (eds), The University of Texas at Austin, 169-176: 1998[2] B. R. Birmingham and H. L. Marcus, Silicon Carbide Shapes by Selective Area Laser Deposition Vapor Infiltration, Proceedings of the Solid Freeform Fabrication Symposium, H. L. Marcus et al.

Finite Element Mod.el of Multi-Layered Laser Sintered PartsE.M. Weissmanland M. B. Hsu2AbstractA finite element heat transfer analysis is applied to the selective laser sinteringof a layered part madefrompolyroerpowder. The sinteringsubroutine in the code is based on the analyses of Scherer [l t2] and Mackenzie and Shuttleworth [3]. The density and conductivity of the particle bed are treated as functions of the void fraction of the bed. The Yagi - Kunii [4] thermal characterization of the powder bed is used to calcu late the effective conductivity of the bed. An example is worked for ABS powder.IntroductionUpon hcatingt abed of polymeric powder undergoes a temperature rise and a densification. To better understand this process considerable effort has been expended recently to mathematically modelit [5t6]. A successful model would at a minimum provide density (or void fraction) and temperature as a function of time everywhere in the sinteringpowder bed.

A successful model would at a minimum provide density (or void fraction) and temperature as a function of time everywhere in the sinteringpowder bed. Preferably it would also yield the distribution of stresses that can develop in the sinteredpartas it cools.Jnput~ fora candidate model include the spatial and time param eters of the inputfluxt which wouldinclude the path followed by the laser beam in scanning out an arbitrary partt the initialtemperature of the powdert other thermal boundary conditions and the thermo-mechanical material properties of the constituents of the powder bed and sintered material.The problem of modeling the sintering of a bed of particles on a microscopic scale (Le.tsmaller than the particles themselves) is a computationally daunting task. 1ft insteadt the particle bed is conceptualized as a continuum with effective thermal propertiesthat may depend on temperature and porositYt the task is made more tractable.

1ft insteadt the particle bed is conceptualized as a continuum with effective thermal propertiesthat may depend on temperature and porositYt the task is made more tractable. This is the approach used in Ref.S and in·this work:. conductivity and specific heat of a homogenized particle bed are used in the model. The effective specific heat of the bed is the same as that for the solid particles because of the small mass of air.

The effective specific heat of the bed is the same as that for the solid particles because of the small mass of air. Thermal conductivity is calculated from the Yagi - Kunii [4] relationship:where K(eff, !l,g) =thermal conductivities: effectivet solidt and gase = void fraction ~=I <l> =0.034While the particle bed is treated homogeneously with respect to its thermal propertiest the sintering behavior is idealized in another way: the bed is assumed to consist either of a cubic array of cylinders with open porestas first suggested by Scherer[l]t ort if the density is at least 0.94 of the theoretical maximumt of an array of spherest as proposed by Mackenzie and Shuttleworth [3]. With these two idealizations we can account for both the heat transfer and resulting softeningt flow and coalescence of the irradiated material.1. BFGoodrich Research and Developmentt Brecksvillet Ohio 2.

BFGoodrich Research and Developmentt Brecksvillet Ohio 2. Marc Analysis Research Corporationt Palo Altot California86(EQ I)combined approach to amodel (i.e., a I-D thermal/Scherer model) was first used by Sun difference algorithm to follow the sintering process in time and space we have attempted to solve the problem with the finite elementet al[S]. In that work they used a in a single layer of powder. In this analysis for a sintered part of arbitrary geometry. Our first method, with our ultimate goal demonstrate the successful application of this technique to the same prob- steps, which are presented lem solved in Ref. S, and extend that analysis to the case of a multiply layered part. To carry out the finite we have used the MARC finite element package, to which was appended an appropriate element subroutine to include the sintering submodel. Ultimately we believe that the FEM will provide user the necessary made by the ear heat conduction equations.

Ultimately we believe that the FEM will provide user the necessary made by the ear heat conduction equations. Furthermore, MARC is setup to accept user written subroutines.afullto studysintering process in parts of complex geometry, which are routinelyprocess. In particular we have used the MARC code because of its ability to solve nonlin-submodelThe one dimensional form of the heat transferwas used, which is representative of the physicalsituati<:m of a ends of themoving laser heat source scanning over a bed of low thermal diffusivity far from theThe equations are therefore:(EQ2)with the t-nunUJ1nCT initial and boundary CO(n:l1tlOns:I)qo for 0 S; IS; 'tq (1) {O t:> 'tI) ~Ofor z ~ 00,t:> 0(EQ3)0)Towhere: T T(z,t) :::::field withto position and timeK::::: effective bed conductivity heat:::::p:::: density q laser flux laser 'tdurationboundary condition at Z ::::: 0 assumes that all the energy incident on the bed is absorbed.

At a given teltlpel'ature the thermal 2 will change with time due to the sintering process, which causes a reduction in void fraction. In this work only the density and conductivity changes are tracked and updated. An average value for the may also be explicit functions of temperature, this but we haveinheat is used.in this work.

An average value for the may also be explicit functions of temperature, this but we haveinheat is used.in this work. Future extensions of the model will include itThe functional aeJ:lenclem;y of the conductivity on void tral:bClln is given in EQ 1; that of the density p isby:( l - £)(EQ4)wherefun ----.---J.The finiteapproximation to the heat conduction equation 4, is discussed in a later section.tivity1 and modified density2, using effective thermal conduc-87The·· Scherer and Mackenzie...Shuttleworth ModelsTo follow the evolution of the change in density or void fraction as a function of time and temperature, the Scherer model was used for the low density phase of the process (p S 0.94 PS>, and the Mackenzie - Shut tleworth model was used for higher densities. Both of these models assume that the surface energy reduction of the sintering powder drives the process through viscous mass flow dissipation.

Both of these models assume that the surface energy reduction of the sintering powder drives the process through viscous mass flow dissipation. At the lower densities the powder is assumed to consist of an open pore network of cylinders arranged cubically, with the cylinder diameters equal to the particle diameter and the cylinder lengths proportional to the pore diameter. As sinter ing proceeds, the cylinder heights collapse, reducing the void fraction in the powder bed, until the cylinder walls touch one another. At this point the situation can be described by contiguous spheres with closed pores. The essential relationships for our needs that follow from this approach are:ae at = - 11M (31t) 6V32 - 3ex Vx (1- ex) 2for open pore, low density powder beds and.partially sintered parts(Le. p S 0.94ps)' andae Ml 41t - = - ........ - (--) at 112 31/3Ps // (- ..... 1) P213a;of the sintering material.

Inpnnciple,however,theratioM!rlcanbe Qbtaineddirectly from density vs. time measurements, as wa.spointedout by Scherer[1].The·strainratefQrlQwdensity smterlng,EQ.5,isalSQseentobe afunctiQn·ofx'i Thisis theiaspect ratio Qfthecylin(lersthatmake.uPitheunitc.ellinthatrnotJel.•• (The cQnstant cin ·that.equatiQnis8...J2131t.) In. the Mackenzie-/Shuttleworthmodel, EQ6, strain rate isa fUnction of the ratio Qf the densities oCthe fully sintered materlalpsand the actualdensity ofthe bed p.As the particles smter, thevQidfraction e>drQps.

The Scherer model gives anexpressiQn for e as a function of the cylinder aspectratio in the unit cell as:2 e = 1 - 31tx + 8",,2x = 1{;\ 3Given a value of the particle bed density or void fraction, EQ 9 Can be used to find the correspondingvalue oCx whichin·turn can be substituted intoEQ 5 to find a new strain increment Ae and a new Ae.88(EQ5)(EQ6)(EQ7)(EQ8)(EQ9)ei +When the subroutine is calledt the viscosity 1l(T) is found by EQ 8. If the density is greaterthan 0.94 of full densityt the Mackenzie - Shuttleworth model holds and EQ6 is used to finditheupdated free strain.

numerical models. A discussion. on the technical approach is given first and then results of single and multi-layer sintering processes are presented.Technical ApproachIn the finite element formulation the discretization of both the space and the time variables serves as the basis of the approximation.

The region of interest s in the sinteringmaterial is subdivided into elements, .ilnd the temperature T(s) within an element is interpolated from the nOdal values T ofthe element through the interpolation functions N(s),T(s) = N(s) TThe governing equation of the heat conduction problem can be obtained from the.minimization ofweighted residuals of the differential equation,resulting in the matrix equation :C(T)t+K(T)T = Qwhere C(T) and K(T) arethe temperature dependent heat capacity and thermal conductivity matrices, T is the nOdal temperature vector,t is the time derivative ofthe temperature vectoftand Q is the heat flux vector.The system matrices C(T) and K(T) are obtained from the summation ofall element matricesCe(T) andKe(T) in the mesh asC (T) =LCe (T)K (T) = LKe (T)The element matrices Ce(T) and Ke(T) can be evaluated from the following volume integralsoverthevolume ofthe element:Ce = JNTpcNdVvKeJBTkBdV v89(EQ 10)(EQ 11)(EQ 12)(EQ 13)(EQ 14)In EQ 14 P is the mass density, c is the specific heat and k is the thennal conductivity of the material.The matrices N and B are the shape function and temperature gradient, respectively.

As discussed in the previous section, the sintering material can be characterized by effective conductivity keff and density p, which are dependent on the void fraction e. It is obvious that once keff and p are given then both the heat capacity and the conductivity matrices in the finite element equation can be easily calculated:Ce == JNTp (e) cNdVvJBTke//(E)BdV vNote thatcandkeffin the above equation can also be dependent on temperature.AlIDoughiIDetimevaria.blecan.be discretize4the same.way as space variables, a finite difference schemeis usuaIly chosenin the finite element analysis. The selection of a backward difference scheme yields thefollowing\expression:[~t C (T) +K(T) JTn ==+(T)Equation 16 computes nodal temperatures for each time increment At. It can be seen that starting fromthe initial conditions the. temperatureiscalculatedat aninitial time step for aU elements.

It can be seen that starting fromthe initial conditions the. temperatureiscalculatedat aninitial time step for aU elements. The effective conductivity, density (void fraction) and free strain are updated in the sintering subroutine. The new conduc tivitiesiand densities are thenused to find the temperatures atthe next time increment.During this investigationtwo finite element heatconduction analyses were carried out for validating thepreviously·discussed sinteringandbeattransfer.models.•The.twoianalyses.simulate layered sinteringprocesses, respectively. Both the 2-nodethree..dimensional heat element and the 4-node two--dimensionalplanarheatelement were used for the analyses and identical results were observed. The numerical results presented in this section are those of the 4-node model.As shown in Fig. 1 the single-layered model consists of 304-node elements and 62 nodes. A concentnlted laserfluxqo was applied atone end (nodes 31, 62) ofthe modelfor a duration of1: seconds.

1 the single-layered model consists of 304-node elements and 62 nodes. A concentnlted laserfluxqo was applied atone end (nodes 31, 62) ofthe modelfor a duration of1: seconds. Both the initialandbounctaryconditionsan4IDematerial>properties were obtained from Ref. 7 (Table 1). Figures 2 and 3 show time variationsofapplie<1laser flux and temperature at the pointofapplication (node 31). The void fraction distribution along the depth of the model at 0.008 seconds is depicted in Fig. 4. The character istics of the sintering process can be clearly seen from these results.The multi-layered model iSisimilar to the single-layered model but has 90 elements for three layers. Thesallle concentratedJaser flux was applied sequentially in •three passes to the end of the submodel for each layer(nodes 31 and 122 for the firstlayer, nodes 61 and 152 for the secondlayer and nodes 91 and 182 for the third layer). The duration of each flux application was the same 1: seconds.

The duration of each flux application was the same 1: seconds. Both the initial/boundary conditions and material properties were the same as in the single..layered model.In orderto deal with the mUlti..layeredsituationtheprogram options ACTIVATE andPEACTIVATE were usedfor the representation ofIDe first pass (elements 1 to 30),the second pass (elements 1 to 60) and the third pass (elements Ito 90)ofthesinteringprocess. Although the total number ofelements in the model is90,with the ACTIVATE and DEACTIVATE options there are actually 30 active elements in the first pass, 60 elements in the second pass andi90 elements in the third pass analyses, respectively.90(EQ 15)(EQ 16)Figure 5 shows a finite element model using the 4-node planar heat transfer element for the layered model. Variations in time of laser flux and temperature are depicted in Figs. 6 and 7, respectively. Note that a transient time. of 0.02 seconds is assumed. for each pass.

Variations in time of laser flux and temperature are depicted in Figs. 6 and 7, respectively. Note that a transient time. of 0.02 seconds is assumed. for each pass. FiguresS,9,10 and 11,12, and 13show distributions along the layer depth of temperature and void fraction at transient times of 0.02,0.04 and 0.06 seconds, respectively.The finite element results presented in this section are reasonable. However, a panunetric study wouldbe very useful in determining the important parametersin the process. The intensity and duration of laser flux, temperature dependency of material propetties,as well as thermal<andmechanical boundary conditions are of great importance in the process.TABLE 1. Initial conditions, boundary conditions and material properties (ABS) [Ref.

Initial conditions, boundary conditions and material properties (ABS) [Ref. 7]Boundary Conditions:Laser flux, qo Laser duration, t6.6 x 107 W/m2 4.76 x 10-4 sInitial Conditions:Initial bed temperature, To Initial density, PoMaterial constants:Full Density, Ps Viscosity coefficient, A Activation energy, AE. Surface energy, 'Y Particle bed specific heat, Cp Conductivity of solid ABS, Ks Conductivity of air, Kg1095 Kg/m3 5.41 x 10-18 Pa·s 20638 OK 45 x 10-5 J/m2 1580 JrK.Kg0.21WfK·m 0.026WrK·mConclusionWe have presented the first results of efforts to applythefiniteielement method to the problem of laser sintering. This has been accomplished byincorpor~tingasubroutine ·for the. sinteringmodels/ofSCherer.andable to predictnot only thebchavior in a single layerof material, butalso how abcdofpowder sinters that is laid down in discrete layers in time.

sinteringmodels/ofSCherer.andable to predictnot only thebchavior in a single layerof material, butalso how abcdofpowder sinters that is laid down in discrete layers in time. Our next efforts will focus on extending the model to two dimensions in orderto assess edge effccl~.91References1. Scherer, G., "Sintering of Low·Density Glasses: I, Theory," J. Am. Ceramic Soc., 60, [5· 6], pp 236· 239,1977.2. Scherer, G., "ViscousSinteriragundera.Uniaxial Load," J. Am. CemnllC Soc., 69,[9], pp 206·207, 1986.3. MackenZie,J.,and~b\lttleworth,R.., "Phenomenological Theory of Sintering," Proc. Phys. Soc. London, 62, [12·B], pp833·852, 1949.4. Yagi,S., and Kunii, D., 4'Studieson Effective Thermal Conductivities ira Packed Beds," A.I.Ch.E. Jour· na.I,3. 3,pp 373 • 381, 19575. S\lJl,M.,Bea.man,J.,andB~Iow,J.,''PanpnetricAn~ysisofthe Selective Laser Sintering Process," Prt:>c.SoUdFreeform FabricarlonSym.,pp146 ... 154, August 1990.6.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFINITE ELEMENT MODELING OF HYBRID ADDITIVE MANUFACTURING BY LASER SHOCK PEENINGM.P. Sealya, G. Madireddya, C. Lib, Y.B. Guoba Dept. of Mechanical and Materials Eng., University of Nebraska, Lincoln, NE 68588, U.S.A. b Dept. of Mechanical Eng., The University of Alabama, Tuscaloosa, AL 35401, U.S.A.AbstractHybrid manufacturing has traditionally targeted efficiency and productivity as improvement criteria. However, the advent of additive manufacturing to print functional parts has expanded the possibilities for a hybrid approach in this field. Hybrid additive manufacturing is the combination of two or more manufacturing processes or materials that synergistically affect the quality and performance of a printed part.

Hybrid additive manufacturing is the combination of two or more manufacturing processes or materials that synergistically affect the quality and performance of a printed part. Hybrid additive manufacturing allows for advancements in material properties beyond efficiency and productivity. Mechanical, physical, and chemical properties can be designed and printed. The purpose of this study was to model a hybrid additive manufacturing process to investigate the resulting mechanical properties. Laser shock peening (LSP) was coupled with selective laser melting in a 2D finite element simulation in Abaqus to quantify the resulting residual stress fields. The effects of peak pressure and layer thickness were studied when coupling laser shock peening with selective laser melting.Keywords: hybrid, SLM, laser shock peening, Ti64, residual stress1.

The effects of peak pressure and layer thickness were studied when coupling laser shock peening with selective laser melting.Keywords: hybrid, SLM, laser shock peening, Ti64, residual stress1. IntroductionIn today’s manufacturing industry, the continuous requirement of high quality materials with preferential mechanical properties requires the use of various manufacturing techniques. Additive manufacturing (AM) is a promising method to avoid several limitations of conventional manufacturing processes; however, AM still presents certain challenges in terms of manufacturing high quality parts. For example, thermal loads in metal AM cause tensile residual stresses [1] which can be detrimental to fatigue. Furthermore, surface finish is poor in some AM processes such as directed energy deposition. An alternative to conventional AM is hybrid-AM. A hybrid approach combines two or more manufacturing processes that synergistically affect the quality and performance of a part [2].

An alternative to conventional AM is hybrid-AM. A hybrid approach combines two or more manufacturing processes that synergistically affect the quality and performance of a part [2]. The quality is superior to that achieved by an individual process. A common example of hybrid-AM combines additive manufacturing with subtractive machining [3-12]. In doing so, a workpiece attains a superior surface finish during printing on internal and external surfaces. Other areas of hybrid-AM laser re- processing[17-19].include rolling [13-16] andHybrid additive manufacturing is an emerging area of research. Sreenathbabu and Karunakaran investigated hybrid adaptive layer manufacturing which incorporates printing, heat treating, and machining each layer for rapid tool production [20]. Heat treating relieves stress and strengthens the tool. After the heat treatment, the layer is machined from near net shape to the final dimensions.306Shiomi et al.

Heat treating relieves stress and strengthens the tool. After the heat treatment, the layer is machined from near net shape to the final dimensions.306Shiomi et al. used selective laser melting followed by heat treatment [1]. Strain gages were used to calculated residual stresses after SLM. They observed that large tensile stresses remain on the surface after SLM and found that heat treating reduced the tensile stresses by approximately 70%. Qian et al. proposed hybrid plasma-laser deposition manufacturing (PLDM) which combines the heat sources of the plasma arc beam and laser [21]. They observed that the physical properties of the surface coating were better than plasma deposition manufacturing alone. To the authors knowledge, the only known hybrid-AM simulation was by Zhou et al. [16]. They simulated the microstructural evolution from a hybrid deposition and micro-rolling process to study the effect of rolling on dynamic recrystallization.

[16]. They simulated the microstructural evolution from a hybrid deposition and micro-rolling process to study the effect of rolling on dynamic recrystallization. Others have investigated hybrid welding processes [22-24].The objective of this paper is to propose a new manufacturing process which combines laser shock peening (LSP) and selective laser melting for hybrid additive manufacturing. LSP is a surface treatment used to improve the mechanical properties of a material. The laser irradiates a target and produces a shock wave that plastically deforms the workpiece. This shock wave hardens and strengthens the surface. The effects of LSP can extend several millimeters below the surface.Many researchers have investigated the effect of LSP on different materials and confirmed that LSP can improve the properties by inducing compressive residual stresses [25-28].

Additive manufacturing is a process of joining materials to make parts from a 3D model, usually layer upon layer, as opposed to subtractive manufacturing and formative manufacturing methodologies. Additive manufacturing processes, such as selective laser melting (SLM), selective laser sintering (SLS), and electron beam melting (EBM), can produce products with tensile residual stresses [1,29]. It is hypothesized that these tensile stresses can be altered by employing laser shock peening between printed layers. With the change in residual stresses and increased hardness, the performance of a part can be improved or manipulated. This hybrid additive manufacturing approach enables one to design and print preferential mechanical properties for specific applications. In this study, hybrid additive manufacturing via laser shock peening was modeled in Abaqus. The effects of layer thickness and peening pressure on the residual stresses were studied.2.

In this study, hybrid additive manufacturing via laser shock peening was modeled in Abaqus. The effects of layer thickness and peening pressure on the residual stresses were studied.2. Finite Element Model2.1 Model DesignA series of two-dimensional finite element models were developed in Abaqus Standard to simulate consecutive layers of selective laser melting (SLM) followed by laser shock peening (LSP) on Ti64. The simulation was based on the work from Sealy and Guo [30,31] and Chao and Guo [32]. The objective was to model a hybrid additive manufacturing process, i.e. SLM and LSP, to determine the effect of successive printed layers on the enhanced mechanical properties from LSP. The simulation procedure applied a moving heat flux in a thermal model and imported the resulting temperatures into a stress model. Importing temperatures accounted for thermal strains caused by the heat flux.

Importing temperatures accounted for thermal strains caused by the heat flux. After allowing the temperatures to cool for 5 seconds, laser shock peening was applied to each printed layer via a shock pressure load. This process was repeated for each printed layer.307The workpiece was divided into four parts: substrate, layer 1, layer 2, and layer 3, see Fig. 1. The substrate was 4 mm (length) by 1 mm (thickness). Each layer was 2 mm long. Three different layer thicknesses were investigated: 100 µm, 300 µm, and 600 µm. In each model, three layers were printed and subsequently peened. Layer build-up was accomplished by first deactivating the entire mesh and then activating each layer in each active heat flux step. The thermal model used 4-node linear diffusive heat transfer elements (DC2D4), and the stress model used 4-node bilinear plain stress elements with reduced integration (CPS4R).

The thermal model used 4-node linear diffusive heat transfer elements (DC2D4), and the stress model used 4-node bilinear plain stress elements with reduced integration (CPS4R). Plain stress elements were chosen as one single printed line was assumed to be a thin body.(2)Laser peening pressure (static)vHeat flux (moving)Case 1: 100 µm Case 2: 300 µm Case 3: 600 µmlayer 3 layer 2 layer 1(1)Line A1 mm2 mmSubstrate4 mmFig. 1 Two-dimensional model of hybrid additive manufacturing.2.2 Material ModelThe material used for this model was Ti64. The physical and thermal properties as well as the temperature dependent thermal and mechanical properties used in the analysis are shown in Tables 1-3 [33-36]. At room temperature, the elastic modulus was 110 GPa and the yield strength was 910 MPa. At 1655 °C, the elastic modulus and yield strength decreased to 10% of that at room temperature.

At room temperature, the elastic modulus was 110 GPa and the yield strength was 910 MPa. At 1655 °C, the elastic modulus and yield strength decreased to 10% of that at room temperature. Since laser peening involves high strain rates (on the order of 106) that significantly affects the flow stress, future studies will incorporate more complex material models such as Johnson-Cook or an Internal State Variable (ISV) plasticity model based on the BCJ model to capture such rate dependent effects.Table 1 Physical and Thermal Properties of Ti64Density Latent heat Solidus Temp Liquidus Temp(kg/m3) (J/kg) (°C) (°C)4428 365200 1605 1655308Table 2 Temperature Dependent Thermal Properties of Ti64Temp.

The output of the heat flux (F) as a function of position (x) and time (t) was given by the following:2−2(𝑥𝑥−𝑣𝑣𝑣𝑣) 𝑟𝑟2𝐶𝐶𝐶𝐶 𝜋𝜋𝑟𝑟2 𝑒𝑒 where C was the absorption factor, P was laser power in watts, r was the laser beam radius in meters, and v was the scanning speed of the heat flux [37]. The process parameters for the applied heat flux during selective laser melting are given in Table 4.𝐹𝐹 =Table 4 Process Parameters for Applied Heat Flux during Selective Laser MeltingLaser power Scan speed Layer thickness Laser spot radius Scan length(W) 20(mm/s) 50(µm) 100, 300, 600(µm) 36(mm) 2309(1)Stress Model: Laser peening was applied in the stress analysis after selective laser melting. To model laser shock peening, a simplified pressure load was applied to capture the highly transient, dynamic nature of a shock wave produced by plasma expansion. The laser spot size from peening was 500 µm.

The laser spot size from peening was 500 µm. The pressure pulse was assumed to be 2-3 times longer than a typical 5-7 ns laser pulse [38,39]. The pressure pulse width was 20 ns, and the peak pressure was 1 GPa or 2 GPa. The peening pressure Prtas a function of both radial position and time was given by(,)2−𝑥𝑥 2 2𝑟𝑟𝐶𝐶(𝑟𝑟, 𝑡𝑡) = 𝐶𝐶(𝑡𝑡)𝑒𝑒represents pressure at any time t, r is the radius of the laser spot [40]. Typically, the where P pressure pulse as a function of time has a Gaussian profile with a short rise time [38,41-43]. In this study, a simplified triangular pulse of the pressure as a function of time was used (Fig. 2). The pressure was applied in the center of the mesh.t()P (GPa)PmaxShort rise time (SRT) Pressure PulseTriangular pulse0td2tdTime DurationFig. 2 Pressure-time history of a single pressure pulse from LSP.2.4 Initial and Boundary ConditionsThe initial temperature of the model was 20°C.

2 Pressure-time history of a single pressure pulse from LSP.2.4 Initial and Boundary ConditionsThe initial temperature of the model was 20°C. During the stress analysis, the nodal temperatures were imported from the thermal model as a prescribed condition during the active heat flux step. Heat was allowed to conduct through the material. No heat transfer boundary conditions were prescribed. The displacement and rotation degrees of freedom along the bottom of the substrate were constrained in the analysis.3. Results and DiscussionThe temperature and stress fields are plotted along the depth direction at the center of the mesh, i.e. along Line A in Fig. 1. The temperature profiles are shown when the heat flux was directly over Line A.

along Line A in Fig. 1. The temperature profiles are shown when the heat flux was directly over Line A. For the stress profiles, the heat flux passed over the layer and the stress S11 was plotted 0.1 ms after the pressure pulse.310(2)3.1 DisplacementThe magnitude of displacement (U) for a 100 µm, 300 µm, and 600 µm layer thickness exposed to a 1 GPa and 2 GPa peening pressure is shown in Fig. 3. It was observed that 1 GPa peening pressure did not cause severe plastic deformation. The maximum transient deformation was approximately 10 µm during peening. The maximum deformation after relaxation was approximately 2-3 µm. At 2GPa, the deformation was considerably higher. The deformation following relaxation was greater than 200 µm. Also, it was observed that deformation in 300 µm layer model was higher than in the 600 µm layer model.

The deformation following relaxation was greater than 200 µm. Also, it was observed that deformation in 300 µm layer model was higher than in the 600 µm layer model. This may be attributed to the expansion during the thermal load varies depending on layer thickness and affects the deformation.e r u s s e r P P S La P G 1U, Magnitude (m)e r u s s e r P P S La P G 2Layer thickness 100 µmLayer thickness 300 µmLayer thickness 600 µmFig. 3 Magnitude of deformation in a 100 µm, 300 µm, and 600 µm layer thickness model exposed to a 1 GPa and 2 GPa peening pressure.3.2 Nodal TemperaturesNodal temperatures from the applied heat flux for a 100 µm, 300 µm, and 600 µm layer thickness are shown in Fig. 4. The blue triangles represent the temperatures while printing layer 1, the red squares represent the temperature while printing layer 2, and the black circles represent the temperature while printing layer 3.

In selective laser melting, the top surface is constantly changing by adding subsequent layers. Therefore, note that a zero depth corresponds to the top of layer 3. If the layer thickness is 100 µm (Fig. 4a), the top of layer 1 corresponds to a depth of 200 µm. For a 600 µm layer thickness (Fig. 4c), the top of layer 1 corresponds to a depth of 1200 µm.The nodal temperatures exceeded 3000 °C on the top surface. For the given heat flux conditions, the results indicated that when the layer thickness was 100 µm (Fig. 4a), the temperature in layer 1 while printing layer 2 was between 800 °C and 1800 °C. The temperature311was relatively high considering that the solidus temperature of Ti64 was 1605 °C. This indicates that part of layer 1 was re-melted during the printing of layer 2.

The temperature311was relatively high considering that the solidus temperature of Ti64 was 1605 °C. This indicates that part of layer 1 was re-melted during the printing of layer 2. When printing layer 3, the temperature in layer 1 ranged from 200 °C to 600 °C.Nodal Temperature (°C) 0 20004000Nodal Temperature (°C) 0 2000 4000Nodal Temperature (°C) 0 2000 40000250layer 3 layer 2 layer 10500layer 3layer 20layer 3)m µ ( h t p e D500750(a)Heat L1 Heat L2 Heat L3)m µ ( h t p e D10001500(b)layer 1Heat L1 Heat L2 Heat L3)m µ ( h t p e D10002000(c)layer 2layer 1Heat L1 Heat L2 Heat L3100020003000Fig. 4 Nodal temperatures (NT11) along the depth direction below applied heat flux for a (a) 100 µm, (b) 300 µm, and (c) 600 µm layer thickness.When the layer thickness was 300 µm, the maximum temperature in layer 1 while printing layer 2 was approximately 450 °C.

The maximum temperature in layer 1 while printing layer 3 was 75 °C. As the layer thickness increased to 600 µm, the temperatures in layer 1 while printing layer 2 was nearly room temperature. The results indicated that for the given heat flux in this study, the layer thickness needs to be greater than 300 µm to avoid significantly raising the temperature of previously laser peened layers. Models such as this can help determine the critical layer thickness for hybrid additive manufacturing processes so that thermal loads from SLM do not negate enhanced mechanical or physical properties in prior layers.4.3 Residual Stress after LSPThe residual stress (S11) along the depth direction after laser shock peening with a pressure of 1 GPa (Fig. 5) and 2 GPa (Fig. 6) with a layer thickness of 100 µm, 300 µm, and 600 µm is shown below. For a layer thickness of 100 µm, 1 GPa was not significant enough to cause compressive residual stress in the printed layer, see Fig. 5a.

For a layer thickness of 100 µm, 1 GPa was not significant enough to cause compressive residual stress in the printed layer, see Fig. 5a. The stress in each layer was a tensile and between 400 MPa and 800 MPa. When the pressure increased to 2 GPa (Fig. 6a), significant compressive residual stresses were imparted in layers 1, 2, and 3. When printing subsequent layers, e.g. layer 3, the heat flux caused the compressive stress from peening in layers 1 and 2 to turn tensile. The layer thickness was relatively small such that the thermal load from the subsequent layer’s heat flux may be expanding the workpiece such that the compressive residual stresses are reversed.

The layer thickness was relatively small such that the thermal load from the subsequent layer’s heat flux may be expanding the workpiece such that the compressive residual stresses are reversed. Furthermore, the higher pressure (2 GPa) coupled with such a thin layer resulted in layer 1 shifting to compression when layer 3 was peened.312Stress S11 (MPa) 0 5001000 -5001000Stress S11 (MPa) 0 5001000 -5001000Stress S11 (MPa) 0 5001000 -5001000layer 30layer 300layer 2 layer 1250layer 2500layer 3)m µ ( h t p e D500)m µ ( h t p e Dlayer 11000)m µ ( h t p e D1000layer 2layer 12000(a)750Post-LSP L3 Post-LSP L2 Post-LSP L1(b)1500Post-LSP L3 Post-LSP L2 Post-LSP L1(c)Post-LSP L3 Post-LSP L2 Post-LSP L1100020003000Fig.

5 S11 along the depth direction after laser peening with a pressure of 1 GPa for (a) 100 µm, (b) 300 µm, and (c) 600 µm layer thickness.Stress S11 (MPa) 0 5001000 -5001000Stress S11 (MPa) 0 5001000 -5001000Stress S11 (MPa) 0 5001000 -5001000000layer 3layer 2layer 3500layer 3250layer 1500layer 2)m µ ( h t p e D500)m µ ( h t p e D1000layer 1)m µ ( h t p e D10001500layer 2layer 12000(a)750Post-LSP L3 Post-LSP L2 Post- LSP L1(b)1500Post-LSP L3 Post-LSP L2 Post-LSP L1(c)2500Post-LSP L3 Post-LSP L2 Post-LSP L1100020003000Fig. 6 S11 along the depth direction after laser peening with a pressure of 2 GPa for (a) 100 µm, (b) 300 µm, and (c) 600 µm layer thickness.In the 300 µm layer thick model, a similar trend was observed.

The high tensile stresses in between the layers were due to the fact that the heat flux from a subsequent layer caused significant thermal expansion and resulted in high tensile stresses. With 2 GPa peening pressures, the maximum compressive stress was more than 900 MPa. The tensile stresses in layer 1 developed during thermal loading of layer 2 remained tensile even after the peening layer 2.Similarly, in the 600 µm layer thick model, the tensile stresses joining two layers were due to the thermal loads and the compressive stresses within the layers were due to peening. The compressive stresses were not as high as those in the 300 µm layer thick model because all of the313heat energy accumulated in one thick layer. Significant thermal expansion ensued and led to tensile stresses on the order of 500 MPa. Subsequent peening did not have a significant effect because the layers were much thicker.The maximum tensile stresses occurred at the boundary between layers.

Subsequent peening did not have a significant effect because the layers were much thicker.The maximum tensile stresses occurred at the boundary between layers. This was due to an excessive amount of thermal expansion during thermal loading and the fact that layers were added as plates on one another. In reality, the starting material in SLM is powder which would be deposited and melted to the substrate. In that case, there should be less thermal expansion than what was observed in a plate.5. Summary and ConclusionsThis work presents a finite element model of a hybrid additive manufacturing process that couples selective laser melting (SLM) and laser shock peening (LSP). The model adds a new layer that is subsequently laser peened. The process is repeated for three layers. The objective was to quantify the effects of single shot LSP after printing multiple layers.

The model adds a new layer that is subsequently laser peened. The process is repeated for three layers. The objective was to quantify the effects of single shot LSP after printing multiple layers. In other words, how does the thermal load from subsequent printed layers influence the residual stress imparted by LSP in prior layers? The effects of laser peening pressure and layer thickness on the residual stress fields were analyzed. This model established the fact that layer thickness plays a critical role on the resulting residual stress fields. If a layer is thicker than a critical value, the thermal loads from printing more layers will not significantly alter the residual stress field in previous layers. For thinner layers, more substantial pressures are needed to cause deeper compressive residual stresses. The results suggest that peening conditions can eventually be optimized to have the desired residual stress contour for a given application.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFINITE ELEMENT MODELING OF METAL LATTICES USING COMMERCIAL FEA PLATFORMSE. Arrieta *, Jorge Mireles*, C. Stewart*, C. Carrasco*, and R. Wicker*The University of Texas at El Paso, El Paso, TX, 79968AbstractThe introduction of geometrical features into standard solids result in cellular materials with unique performances. The deformation mechanisms originated by the introduced geometry may not be entirely captured by the current commercial FEM software; resulting in inaccuracies in predicting the performance of cellular metals. Additionally, the inconsistency of AM material properties will result in material models with uncertainty, thus, contributing to the inaccuracy of simulations.

Additionally, the inconsistency of AM material properties will result in material models with uncertainty, thus, contributing to the inaccuracy of simulations. The present work shows a process for modeling the strength of EBM Ti-6Al-4V lattices structures; starting from the definition of the convenient experiments to generate the data for the development of material models at different orientation and finalizes with the assignment of these material models to the lattice FEMs. MSC Patran/Nastran is used in this work. Experimental results of the compressive strength of lattice structures are compared with those from the FEM utilizing the different material models created from the experiments.Keywords: Lattices structures, Ti-6Al-4V, EBM, FEA, DICIntroductionLattice structures typically are more damage tolerant, displaying more uniform mechanical properties in comparison to stochastic foams [1].

Numerical models suggest that latticed materials are more tolerant to defects compared to foams. By randomly removing struts, the modulus and compressive strength of open-cell foams decreased more rapidly than for lattices. Although varied results are reported, investigations coincide in the proportionality of the elastic modulus to density of open-cell materials [2]; thus in regards to damage sensitivity, the removal of struts reduces the modulus in a faster manner than the thinning of these [3]. Well-designed sandwich plates with cellular lattices in the core showed to withstand larger blast impulses than solid plates of the same material and weight [4]. Other investigations with sandwich panels revealed the post-yield response under dynamic loads to be insensitive to the low velocity range of impact; furthermore, a more localized deformation was observed for high velocities.

In underwater explosion tests, lattices were found to crush in a progressive manner by a subsequent buckling of struts layers, providing insights as to the effects of filling materials in enhancing plastic deformation of the metal [5]. Thus, for the particular interest of energy absorption applications exposed herein, lattices are selected as better alternatives than foams. Additionally, these non-stochastic architectures have been reported to behave within the characteristic parameters for cellular solids defined in [6], such as slenderness, aspect ratio, relative density and modulus among others [7].765Expressed as the rate of the externally applied energy divided by the amount of mass of the material withstanding it, the normalized absorbed energy is the common parameter to evaluate the capability of materials to dissipate energy.

Empirical evidence reports that the bending of struts, the formation and manipulation of multiple shear bands, and microcracking of the constituent material among many other features, also contribute to the absorption of energy [4]. As more applications are developed to use these materials as energy absorbers, structurally efficient members, or a combination of applications, experimental testing of these materials has become a significant component of their development and study. Experimental testing demands valuable resources; numerical simulations ease the bridging from materials science into engineering applications and increase the chances of early experimental success.

Experimental testing demands valuable resources; numerical simulations ease the bridging from materials science into engineering applications and increase the chances of early experimental success. The development of computer models of cellular materials grants us the availability to modify features on their design and explore their response; all in the same CAD model from which the components will be fabricated by AM.In this context, this work presents the development of material data sets and finite element models of different lattice materials for comparison with experiments of actual lattice structures under compression. The CAD models were developed in Solidworks, while Patran and Nastran from MSC Software were utilized for the development of FEM and FEA respectively.Materials and MethodsAll the specimens were fabricated with EBM Ti-6Al-4V in an Arcam A2 system.

The specimens for testing and comparison with their computer models consisted of two types of lattices: hexagonal unit-cells and reentrant hexagonal unit-cell. Each type of lattice was designed at three different densities from varying the cell size (5mm, 6mm and 7mm) and with length, width and height equal to 5 unit-cells in each direction and the microstruts were 0.8mm x 0.8mm (Figure 1). A monolithic plate was fabricated as part of the lattices, for uniform stress distribution. Three specimens of each density were tested and results averaged. The lattices were tested as per ASTM E9 at a constant cross-head speed equivalent to a strain rate of 0.003 mm/mm ∙ min [8]; (Table 1) in addition, relative densities (ρ*) were calculated by weighting the specimens and subtracting the volume of the plates assuming to have the same density as solid ASTM E9 specimens fabricated and discussed later in this work.

The actual lattices and their computer models were compared based on compressive strength.Table 1 Average compressive properties of lattices.Lattice 5mm HEX, 0.8mm strut 6mm HEX, 0.8mm strut 7mm HEX, 0.8mm strut 5mm RE-HEX, 0.8mm strut 6mm RE-HEX, 0.8mm strut 7mm RE-HEX, 0.8mm strutρ* (gr/cm3) 0.800 0.600 0.470 1.130 0.820 0.650Ultimate Strength (kPa) 40,000 21,000 15,000 35,000 21,000 13,000E(GPa) 2.672 0.943 0.795 1.500 1.042 0.786Specific Strength (kN*m/kg) 50.000 35.000 31.915 30.973 25.610 20.000Specific Modulus E6(m2/s2) 3.340 1.572 1.691 1.327 1.271 1.209766The elastoplastic constituent material models were created from stress-strain curves and mechanical properties obtained from testing microstruts and ASTM standards solid specimens printed at different orientations (Figure 2); strain measurements of the specimens were performed by Digital Image Correlation (DIC).

Experiments showed similar responses for solid Ti-6Al-4V under compression (ASTM E9) and under tension (ASTM E8). In this manner, the material data set consisted of stress-strain curves and mechanical properties measured on ASTM specimens built at 0º, 30º, 60º and 90º [9], in addition to curves and properties measured on individual microstruts, which form the lattices, built at 0º, 30º, 45 º, 60º and 90º. In this context, the material data set permits assigning material properties to finite elements oriented accordingly; for example, the material models generated from testing 30º specimens (ASTM or microstruts) were assigned only to the finite elements in the 30º microstruts; while 90º properties were assigned to the finite elements in the 90º microstruts, in the computer model.

The lattices in this study had microstruts oriented at 90º and 30º only, while the bearing plates were modeled with material properties from ASTM E9 on 90º specimens.Figure 1 As-fabricated hexagonal (left) and reentrant hexagonal lattices at different densities.Figure 2 EBM Ti-6Al-4V specimens for testing and determination of material properties.767Finite Element ModelThe rapid prototyping process started with the 3D-solid CAD designing of the proposed lattices in Solidworks from Dassault Systèmes. This CAD served as the root file, from where the stereolithography (.stl) and the parasolid (.x_t) format files were exported for fabrication and FEM in the ARCAM A2 system and Patran software respectively. The FEM started with importing the .x_t file into MSC Patran, at the correct scale and required tolerance of dimensions (0.01).

The FEM started with importing the .x_t file into MSC Patran, at the correct scale and required tolerance of dimensions (0.01). Three different conditions were observed in the model: solid bearing plates fabricated at 90°, and microstruts fabricated at 90° and 30°; therefore, three different materials were created in the FEM. The load distribution plates were defined as bulk Ti-6Al-4V, fabricated at 90° whose mechanical properties from the experiments were determined as isotropic with E=113GPa and υ = 0.33. The vertical microstruts were defined as 90° microstrut Ti-6Al-4V, whose mechanical properties from the tensile experiments on microstruts at 90° were determined as isotropic with E=113GPa and υ = 0.33. The inclined microstruts were defined as 30° microstrut Ti-6Al-4V, whose mechanical properties from the tensile experiments on microstruts at 30° were determined as isotropic with E=104GPa and υ = 0.31.

The linear elastic and elastoplastic constitutive models were set for both microstrut Ti-6Al-4V properties, whereas for the bulk Ti-6Al-4V only the linear elastic model was set, given the expected strain distribution throughout the lattice. The elastoplastic constitutive models were defined by data fields extracted from averaging the stress-strain curves from the actual experiments (Figure 3). Additionally, the lattices were modeled with material properties from the ASTM E8 standard tension tests, and from the tension test on microstruts.768Figure 3 Stress-strain data field used in the material numerical models of microstrut properties (top) and ASTM properties (bottom). These data fields were discretized from actual measured curves. For clarity, only 30° and 90° ASTM E8 are shown.

These data fields were discretized from actual measured curves. For clarity, only 30° and 90° ASTM E8 are shown. Series names indicate the build angle and last digit as a unique specimen IDDefining T1, T2, and T3 as translations along orthogonal axes 1, 2 and 3, also named x, y and z, where x-z denotes the horizontal plane, the nodes in the bottom surface of the solid model were constrained for z displacements; similarly, for rotations around the axes R1, R2 and R3, these nodes at the bottom were also constrained for R1 and R3. Total constraint was applied only to nodes of the element at the center of this surface (Figure 4).

Total constraint was applied only to nodes of the element at the center of this surface (Figure 4). The upper surface of the plate at the top was exclusively loaded with the vertical element-uniform displacements, corresponding to the769displacements of the actual plate at the ultimate strengths in the experiments; the displacements were measured by DIC (Table 2).The complex geometry of the 3D CAD solids representing the lattices restricted the element selection to tetrahedrons; based on the nonlinear response observed during testing, the quadratic 10-node tetrahedron (Tet10) was suggested, because the extra node on its sides derives in quadratic shape functions permitting to better capture the nonlinear behavior observed in the testing.For the initial simulations, the mesh was automatically generated by the solids mesher TetMesh by MSC Patran.

The mesh density is controlled by the global edge length parameter; where the longest edge in the geometry, divided by this parameter defines the number of elements along the edge. TetMesh recognizes the longest edge from the geometry of the solid, automatically determining the global edge length based on the minimum element edge length, which has a default value of 20% of the global edge length, producing elements of acceptable quality. Thus, the global edge length was different for each size of lattice in meshes generated automatically. As a gauge, that facilitated the measurement of the reaction force of the whole lattice at a single point, an additional multipoint constrain element (MPC-RBE2) collecting all the reaction forces from the surface nodes on the exposed face of the bottom bearing plate was included.

The RBE2 is characterized by a relationship purely dependent on nodal displacements, meaning it does not possess neither stiffness, mass, force, nor any other relationship (Figure 4Error! Reference source not found.).Figure 4 FEM of the hexagonal, 6mm lattice loaded with 1mm element-uniform displacement. MPC and resultant force vector at the bottom.770Finite Element AnalysisThe simulations were run in a Dell Precision T5500 computer equipped with Intel Xeon E5603 processor and 6GB DDR3 SDRAM operating at 64-bit. For the solution, the model was divided into four domains with similar number of elements each, and the maximum RAM compromised to solve the model was set at 80%, saving the remaining 20% for the basic operation of the system. The project is solved in Nastran SOL600; a solver for implicit nonlinear solution, following full Newton-Raphson methods with a maximum of 20 iterations per increment, residual force of 0.1 and automatic abort if failed to converge.

The adaptive feature is set for defining the loading increments.Convergence AnalysisConvergence analyses were conducted for all models to ensure that the mesh density was not considerably affecting the results. Four different mesh densities were explored for each of the models representing each of the lattices. One of the densities was the one produced by TetMesh. Given that the mesher takes into consideration the longest edge in the geometry, the mesh refinement controlled by the global edge length (longest geometrical/number of elements) would prioritize the bearing plates.

A more effective way to properly size the mesh in these solids was found to define and control the separation of mesh seeds along the edges of the microstruts, in other words, adding more elements where the stress is distributed.The convergence analysis compared the resultant force at the RBE2 element, produced from applying the displacements at the ultimate strength in the experiments, with the number of elements in the model. Starting at about 70,000 elements, the convergence curves showed a tendency to consistent results. In regards to the computational work, the models started to be significantly more expensive around the 100,000 elements, where the relative efficiency (number of elements / wall time) escalated with no significant change in the resultant force (Table 2 and Figure 5).

Thus, the mesh seeding along the microstruts was demonstrated as an effective way for the mesh refinement, and that the consistency of the results may be acceptable in the range of 70,000 to 100,000 elements, when the computational work becomes considerably more expensive.771Figure 5 Mesh convergence curves for hexagonal (left) and reentrant hexagonal (right) models. 5mm, 6mm and 7mm unit-cell sizes (top to bottom); two material properties per graph are shown. Solid line representing the experimental average ultimate load.Results and DiscussionThe results of the analyses were evaluated by comparing the nodal force reaction from the independent node at the RBE2 element and the average ultimate force from the experiments.

The models were created in two variants: one using the material properties data from the ASTM E8 tests, and another with the data from the tension tests of microstruts (Figure 3).In general, the models displayed very uniformly distributed stress fields throughout the latticed region, suggesting an efficient energy-distributing material (Figure 6). The models also772indicated that inclined microstruts absorb most of the loading, thus it is advisable to optimize them for an increased capacity (Figure 7). The 90° microstruts showed a lower stress among both configurations, however, the reentrant hexagonal lattices presented columns with lower stresses than the hexagonal. These conditions are the result of the purely compressive state in the hexagonal cells compared to the combination of inclined microstruts in tension and columns under compression in the reentrant hexagonal lattice (Figure 8).

It was observed that, by extending the compressive forces to the columns, the purely compressive state better distributed the stresses to more material, compared to the stress distribution at the nodes in the reentrant hexagonal lattices (Figure 9). This more uniform stress distribution by the hexagonal unit-cells explains the failure at lower loads of the reentrant hexagonal lattices that concentrates most of the stress in the inclined microstruts under tension (Figure 10). In similar manner, the inclined microstruts are observed to support most of the shear stress (Figure 11).

In similar manner, the inclined microstruts are observed to support most of the shear stress (Figure 11). It is important to be aware that, because of the fragility of the Ti-6Al-4V, the reentrant unit-cells failed before completely developing its negative Poisson effect, and before the cells reconfigured for a second, purely compressive, loading stage that might result in stronger materials with the same amount of mass.Figure 6 von Mises stress distribution (MPa) in hexagonal (left) and reentrant hexagonal (right) 5mm lattices. Darker colors indicate lower stress.773Figure 7 Inclined microstruts with higher von Mises stress (MPa).Figure 8 Stress “Y” component in MPa; purely compressive stress state in 7mm hexagonal lattices (left). Tension (-) and compression (+) stresses in 7mm reentrant lattices (left).774Figure 9 Stressed microstruts and low-stress columns in reentrant hexagonal lattices.

On the other hand, results from the reentrant hexagonal lattices showed closer values to the experimental results, when the material properties input data derived from the standard tension tests. On the comparison of material properties, the models with material properties developed from the microstrut tests showed to be more accurate than using data from standard tension tests, this could be explained from the larger sensitivity to manufacturing defects of the microstruts, adding uncertainty; besides the removal of surface defects in ASTM E8 specimens. On the comparison of the selected geometry; the hexagonal geometries exhibited values closer to the experimental results, when compared to the reentrant hexagonal, this could be caused by the more complex stress distribution subjected to tension and compression in the reentrant hexagons, compared to the purely-compressive stress state in the hexagonal lattice.

On the comparison of the size of the unit-cell; the smallest size (5mm), the models showed more accurate predictions; this, given the increased density of the lattice diminishing the geometrical nonlinearities introduced at the mesoscale, in other words, the smaller the unit-cell, the denser the lattice, and the closer of this to uniformly behave as a standard solid.The two unit-cell configurations displayed more uniformly distributed plastic strain fields for larger unit-cell sizes (Figure 12). Another interesting, and expected, feature in the models is that the largest strains developed in the microstuts in the interior of the lattice (Figure 13).

The reentrant hexagonal lattice models were also capable of developing the negative Poisson effect (Figure 14) and measured in [8]; suggesting an additional argument favoring the confidence on the FEM for evaluating the auxetic deformation mechanisms, and therefore the strain field distribution, in other words, the deformation mechanisms in the models were similar to those from the experiments.776Figure 12 von Mises plastic strain hexagonal (top) and reentrant hexagonal (bottom) lattices with 5mm (left) and 7mm (right) unit-cells.Figure 13 Plastic strain in interior microstruts.777Figure 14 Undeformed model wireframe in highlighted light color on the left side face of the lattice, highlighting the auxetic behavior.778s e i t r e p o r pl a i r e t a m d n a y t i s n e dh s e mt n e r e f f i dt asM E F f o y r a m m u S 2 e l b a T. g v A p x E.. v e D d t S.d a o L . t l U. v .o.

g v aN k 0 . 6 1 = F ,m m 8 3 . 0 = . lp s i d (. r r e . g v A. r r e . g v A. r r e . g v A. r r e . g v A. r r e . g v A. r r e . g v A. r r e . g v A)m m 7 ()m m 6 ()m m 5 () X E H E R () X E H() s t u r t s o r c i m() .p o r p.d t s (% 6 0 1 4.% 8 1 5 2.% 3 1 1 2.% 7 4 1 3.% 0 6 6 2.% 2 3 5 2.% 6 7 2 3.ConclusionsAlthough the complex geometry of the lattices limited the selection of elements to tetrahedrons, the resultant compression force in the FEMs show convergence for mesh refinement.TetMesh was found not effective in increasing the density of the mesh for this kind of models. A better approach was presented by mesh seeding the edges of the microstruts.

A better approach was presented by mesh seeding the edges of the microstruts. Following the mesh seeding approach for the mesh generation, the variation of the results, because the size of the mesh, can be neglected for models over 70,000 elements, but no larger than 100,000, when the computational work dramatically increases.On the prediction of the experimental compressive forces applied, some models showed considerable errors (up to 70.65%); however, this error needs to be viewed within the context of the variability in the material properties as shown in [8], where the coefficient of variation of the failure loads for microstruts can be as high 0.353. In general; on the selected geometry, the hexagonal models showed more accurate compression forces compared to the reentrant hexagonal.

In general; on the selected geometry, the hexagonal models showed more accurate compression forces compared to the reentrant hexagonal. On the material properties used for modeling, the models with input data derived from testing microstruts showed more accurate with an average error of 25.32% in predicting the average experimental force, compared to those with input data from ASTM E8 tests, with an average error of 32.76%. On the unit-cell size, the smallest (5mm) showed more accurate in estimating the experimental compression force.

In summary, in predicting the average experimental resultant compression force applied, the model for hexagonal lattices showed to be more accurate, with a 2.99% average error, for the 6mm unit-cell and microstrut properties; whereas the most accurate models for reentrant hexagonal lattices was found with 6mm unit-cell and standard material properties, with an average error of 6.32%.Recommendations and Future Work.On the development of more effective automatic meshers; it is recommended that commercial FEM platforms wishing to start modeling cellular lattices, to allow the user to specify the region of interest where the global edge length, as discussed in the convergence analysis, needs to be defined.On the error of the models predicting the experimental results; it may be advisable to investigate the development of newer technologies for finite elements that allow to capture the nonlinear effects of the geometries introduced at the mesoscale.On improving the accuracy of the presented models, a more elaborated model with mixed elements could be developed.

For example, for mechanistic unit-cells, in regions where high stresses are expected, such as in the nodes of the lattices, elements of higher order are better suited for capturing stress distribution, whereas first order elements may be suitable for the microstruts. However, this could be time consuming, and thus, it might be seen restrictive in rapid prototyping. Although the material properties derived from testing microstruts include the effect of the manufacturing defects, the uncertainty on the properties remains high, because of the size of the microstrut. In that context, and because the nodes of the finite elements are commonly assigned by software in a deterministic manner, the manufacturing defects could also be included in the FEM, by randomly varying the coordinates of the nodes at the free faces of the elements.

The780limits for this random variation could be defined by the metrics of the surface defects in microstruts.Lastly, in further understanding the capabilities and limitations of the commercial FEM/FEM platforms simulating cellular solids with large deformations, it is recommended to evaluate the presented development of these models using more ductile constituent solid such as stainless steel.AcknowledgmentsThis work was supported by the W.M. Keck Center for 3D Innovation, and partially fundedby the NSF grant #1405526.References[1] D. T. Queheillalt and H. N. G. Wadley, “Cellular metal lattices with hollow trusses,” ActaMater., vol. 53, no. 2, pp. 303–313, Jan. 2005.[2] Y. W. Kwon, R. E. Cooke, and C. Park, “Representative unit-cell models for open-cell metal foams with or without elastic filler,” Mater. Sci. Eng. A, vol. 343, no. 1–2, pp. 63–70, Feb. 2003.[3] J. C. Wallach and L. J. Gibson, “Defect sensitivity of a 3D truss material,” Scr. Mater., vol.45, no.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFinite Element Modeling of the Selective Laser Melting Process for Ti-6Al-4VAlaa Olleak, Zhimin XiDepartment of Industrial and Systems Engineering, Rutgers University, 96 Frelinghuysen Rd, Piscataway, NJ 08854AbstractPhysics-based modeling of the selective laser melting (SLM) process is critical for better understanding the influence of the parts quality with respect to various process parameters and scanning strategies. The challenge is to balance model validity, domain the model can be practically useful for size, and computational efficiency so that improving reliability and quality of the printed products. In this paper, a transient thermal finite element model of a SLM process for Ti-6Al-4V is developed using ANSYS for predicting the melt pool size and thermal history.

In this paper, a transient thermal finite element model of a SLM process for Ti-6Al-4V is developed using ANSYS for predicting the melt pool size and thermal history. The thermal solution is the to predict remapped products. The thermo-mechanical model is capable of handling practical domain size with reasonable computational efficiency on the process level, by developing the re- meshing and remapping technique that adapts with the scanning vector.tothe structural problemtheinduced residual stress of1. Introductionthe category of powder bed fusion process. SLM enables the production of metallic parts that have complex geometry without the need of special tooling and multiple processes in a traditional manufacturing process. Parts manufactured by SLM, however, usually exhibit larger product quality variability than the parts produced by the subtractive methods.

Parts manufactured by SLM, however, usually exhibit larger product quality variability than the parts produced by the subtractive methods. In other words, parts quality and reliability control is much more difficult in the SLM, which is mainly caused by different microstructure, induced residual stresses, and manufacturing defects such as high porosity.SLMis a common additive manufacturing (AM) processthat falls underQuality of additively manufactured parts was studied by many researchers and porosity was found to be one of the most important factors that affect parts fatigue takes place due strength and crack propagation resistance. The porosity to either excessive energy density the alloy constituents, or to vaporization of that insufficient energy density that results in incomplete melting of powder [1].

The porosity to either excessive energy density the alloy constituents, or to vaporization of that insufficient energy density that results in incomplete melting of powder [1]. Residual stresses were also found to be important for parts life, which may be influenced by process parameters, building directions and orientations [2], scanning strategy, and temperature of the powder bed [3]. However, the effects of scanning strategy on the residual stresses are not well understood yet [3]. Furthermore, parts microstructure was also found to be influential in terms of the parts strength and life, and it is affected by process parameters and [4].

Furthermore, parts microstructure was also found to be influential in terms of the parts strength and life, and it is affected by process parameters and [4]. Although post-processing techniques including heat treatment, and hot isostatic pressing can help alleviate some aforementioned problems, they cannot eliminate them completely such as the porosity which is just reduced to undetectable range [5].leadsscanningstrategy as well1710Predicting the part thermal history during SLM is important to understand how the part quality and microstructure are influenced by the process parameters. Numerical methods of the SLM process provide useful tools that help understand temperature distribution and history, thermal stresses, and part distortion.

Numerical methods of the SLM process provide useful tools that help understand temperature distribution and history, thermal stresses, and part distortion. Furthermore, they can help optimize the process parameters to achieve better mechanical properties for the printed part as opposed to time-consuming experimental approaches.inducedthe expensive andSeveral studies were developed to simulate the SLM process using finite element models [6–10]. However, the mesh scheme is still a challenge [11]. On one hand, fine mesh the high temperature gradient at the melt pool. On the other hand, a large domain size is necessary to investigate different scanning strategies on the part level and to develop the layer-by-layer model. Therefore, it is technically challenging to have a high-fidelity model for simulating multiple scanning vectors and scanning strategies.

Therefore, it is technically challenging to have a high-fidelity model for simulating multiple scanning vectors and scanning strategies. Hence, some studies had to either reduce the domain size or use large element size compared to the laser spot, which may eventually lead to large discretization error. Furthermore, the mesh configuration and attributes are continuously changing because of the newly added layers and changed material states.isrequiredto accurately capturethe gaussian heat profile andEven though the element death and birth approach was found to be useful to model the addition of the layer, this approach requires large number of elements since the element size is restricted by the layer thickness. Although this approach is used in commercial software packages now, large parts. Another efficient approach is to use adaptive local mesh refinement where the mesh is continuously refined and coarsened according to laser movement [12].

Another efficient approach is to use adaptive local mesh refinement where the mesh is continuously refined and coarsened according to laser movement [12]. This approach time with acceptable accuracy. the computational However, it needs more time in re-meshing. In addition, remapping is computationally expensive in significant frequently. The accumulated errors would accuracy loss at the high temperature gradient regions. Finally, the thermo-mechanical problem is much more expensive because of the higher degrees of freedom at each node. In addition, re-mapping of the mechanical problem needs equilibrium iterations in order for the structure to reach an equilibrium state.itis still computationally expensive forsignificantlyreducesif doneresultIn this paper, a transient nonlinear finite element model is developed to simulate the SLM process. The objective is to handle a large domain size with high accuracy by considering the thermal properties of the thin powder layer.

The objective is to handle a large domain size with high accuracy by considering the thermal properties of the thin powder layer. To reach a compromise between model accuracy and is proposed along with the scanning vector. Basically, fine mesh is employed along with the scanning vector to accurately capture the high temperature gradient with minimal re-meshing steps. The model hence should be able to handle large domain size. After the thermal problem, new mesh with larger element size is created, and the temperature information induced residual stresses in structural analysis. The goal of this study is to predict the temperature distribution and induced residual stresses considering different process conditions and scanning strategies. The model employs Ti-6Al-4V alloy as an example because it is widely used in many applications such as biomedical and aerospace applications in which high reliability and quality are needed.

The model employs Ti-6Al-4V alloy as an example because it is widely used in many applications such as biomedical and aerospace applications in which high reliability and quality are needed. Ti-6Al-4V is a biocompatible materialits computational efficiency, adaptivere-meshingisthen mappedtothe new mesh for predictingthe1711that has high corrosion resistance, high strength at high temperatures, and high strength to weight ratio.2. Model Description2.1 Material modelingA nonlinear material model was used considering material phase changes and material temperature- dependent material models were defined for both the powder and solid materials and they behave thermal conductivity was calculated based on Childs et. al study [13], where the powder thermal conductivity is a function of the thermal conductivity of solid and relative density, equation 1. The volumetric heat capacity for both solid and powder was calculated temperature-dependent density.

The volumetric heat capacity for both solid and powder was calculated temperature-dependent density. The thermal properties are shown in Fig. 1 and the nonlinear mechanical properties are shown in Fig. 2. It is assumed that mechanical and thermal properties are isotropic. The possible changes in the material properties due to the different process parameters are ignored. For the single bead models, the initial temperature was set as 25 oC. For models including scanning strategies, it is assumed that the powder is preheated to 400 oC [14]. For the mechanical problem, the elastic modulus and thermal expansion are set as the very low values after the melting temperature so that the thermal expansion has negligible effect. The material plasticity model is a temperature- dependent multilinear isotropic hardening from the experimental data in [15].transitions,i.e., powder meltingandsolidification. Twosimilarly beyondthe meltingtemperature.

Twosimilarly beyondthe meltingtemperature. The powderfromthetemperature-dependentspecific heat and𝐾" = 𝐾$(𝑇)		(1 − 0.2	𝜀	 − 1.73	𝜀1)(1)(a)(b)Figure 1: Temperature-dependent properties of Ti6Al4V: (a) Volumetric specific heat, and (b) Thermal conductivity [7, 9]2.2 Heat source model and boundary conditionsThe laser heat flux input is represented by a Gaussian heat source expressed in Eq. 2 [11]. In this study, it is assumed that only the elements on the surface are exposed to the laser beam. The heat flux at any node i with coordinates (𝑥3, 𝑦3) can be calculated as a function of time as1712𝑞3(𝑡) = 189 @ :;<=>?𝑒BC1@ DEFGE<=>?(?)H@ IDJFGJ<=>?(?)H @ K<=>?Lwhere a is the absorptivity, P is the laser power, and rspot is the beam radius. The equation parameters are shown in Table 1. As will be discussed later, the effect of to previous absorptivity on simulation is allowed assuming that the surrounding gas is Argon (h=10 W/m2K).

The equation parameters are shown in Table 1. As will be discussed later, the effect of to previous absorptivity on simulation is allowed assuming that the surrounding gas is Argon (h=10 W/m2K). Since the part size is large, the mild steel plate is not included in the model.the resultsis crucial. Two values are used accordingin [17] and experimental resultsin [18]. Convection atthe surfaceFigure 2: Temperature dependent mechanical properties of Ti-6Al-4V2.3 Mesh configurationAccurate representations of the heat source and the melt pool dimension require a small element size in the finite element model. Representing the entire domain using for models small element size, however, considering multiple scanning vectors. Therefore, 8-noded hexahedron elements with edge size of 10 µm were used to represent the powder and melt pool region. For the rest of the domain where the temperature gradient is not high, free 4-noded tetrahedral elements were used.

For the rest of the domain where the temperature gradient is not high, free 4-noded tetrahedral elements were used. For these elements, the material assignment was kept as solid. This the melting temperature is low and the heat flux perpendicular to the laser beam direction is low as well. Fig.3 shows the mesh configuration for a 2 mm scan length with single bead analysis.is computationally prohibitive especiallyis valid onthe groundthat powderthermal conductivity belowTable 1: Process parameters used in the modelPower (Watt) Scanning speeds (mm/s) Absorptivity (a) Spot diameter Layer thickness100,150, and 195 500, 750, 1000, 1200 0.4, 0.6 100 µm 30 µmHaving such small element size allows us to take the powder’s thermal properties into consideration. Therefore, re-meshing is needed to keep assigning small element size where the temperature gradient is high and assigning large element size for the remaining part.

Therefore, re-meshing is needed to keep assigning small element size where the temperature gradient is high and assigning large element size for the remaining part. A typical mesh configurations for a large part (6mm × 6mm × 1mm) for1713(2)both bidirectional and chessboard scanning strategies are shown in Fig. 4. The mesh near the surface always have small size to keep the temperature re-mapping accurate. In the shown mesh in Fig. 4, the size of elements on the top size does not exceed 50 µm.(a)(b)Figure 3: Mesh configuration for single bead model: (a) Block size 3 mm × 3 mm × 0.75 mm, and (b) cross sectional image showing mesh density at melt pool region(a)(b)Figure 4: Mesh configuration for large block size 6 mm × 6 mm × 1 mm: (a) Bidirectional scanning strategy, and (b) Chessboard scanning strategy (Island size: 2 mm x 2 mm).2.4 AnalysisThe analysis is transient in both thermal and mechanical problems.

For the thermal problem, the laser location is updated after each time step, which is the time needed for length). The phase change and the the element edge temperature dependent properties as shown the problem, which requires the full Newton Raphson method. To ensure convergence, the automatic time stepping was turned on in ANSYS. Once the temperature information is fromlaser moving 10 µm (i.e.,in Fig. 1 add high nonlinearitytothe scanning vectoris calculated, new meshis generated, and remapping1714followed to calculate the initial temperature of the new mesh. The thermal solution is calculated from the governing equation shown in Eq.

The thermal solution is calculated from the governing equation shown in Eq. (3).𝜌𝐶"OPOQ= 	𝑘 SOP@OT@ + 	 OP@OV@ + 	 OP@OW@X + 𝑄where 𝜌 is the density, 𝐶" is the specific heat capacity, t is the time, T is the temperature, k is the thermal conductivity, and Q is the heat input per unit volume.Figure 5: Methodology flow diagramThe mechanical problem would be too expensive to solve if keeps the same mesh from the thermal problem with a large stiffness matrix. Therefore, coarse and uniform mesh was preferable. Brick elements with 8 nodes were used for the structural analysis. The nodal temperature as a function of time was calculated from the thermal analysis results and applied to the node as a body force.

The nodal temperature as a function of time was calculated from the thermal analysis results and applied to the node as a body force. The last step of the structural problem is a relaxation step where all the body forces are removed, and the part residual stresses are estimated at the room temperature.Remapping the node temperature from old mesh to new mesh requires locating the nodes from the new mesh in the old mesh, finding the corresponding elements, and then performing element interpolation. However, this process was found to be time consuming. Therefore, a simple weighting function formulated in Eq. (4), was used for into from the consideration for calculating the nodal temperature at the new mesh.remapping.

Therefore, a simple weighting function formulated in Eq. (4), was used for into from the consideration for calculating the nodal temperature at the new mesh.remapping. The nearest 4 nodesthe solved problem weretaken1715(3)𝑇^𝑑^ ≤ 2.5	µ𝑚⎧ ⎪ ∑ ⎨ ⎪ ⎩𝑇3(𝑡) =ef(?) gf j gk where N represents the nearest 4 nodes to node i, Tj is the temperature of node j in old mesh, and dj is the distance from node j in old mesh to node i in the new mesh. The overall analysis flowchart is shown in Fig. 5 with a seamless integration between MATLAB and ANSYS. An example of the re-mapping result is shown in Fig. 6, in which good degree of the computational efficiency.f∈i𝑑^ > 2.5	µ𝑚∑k∈iagreementis maintained while greatlyimproving(a)(b)Figure 6: Temperature remapping Results between different scanning vectors: (a) before remeshing, and (b) after remeshing3. Model ResultsThe single bead model results are compared to the experimental results in [19].

Model ResultsThe single bead model results are compared to the experimental results in [19]. Overall, the melt pool width shows acceptable agreement with the experimental results. However, temperature distribution of a scanning vector using 150 W laser power with 1200 mm/s scanning speed.the melt pool depth was underestimated. Fig. 7 showstheThe effect of the scanning speed and power are compared to the experimental the results the experimental results as shown predicted melt pool width has an increasing trend with the scanning speed and laser power. Though to experimental increases and scanning speed reduces. This deviation of the melt pool depth predictions could be due to the simplification of the heat source model by including only the element on the top surface or ignoring the stochastic nature of the process. Furthermore, the randomly distributed powder particles and their size could affect the powder thermal conductivity and their absorptivity as well.in[12,13].

Furthermore, the randomly distributed powder particles and their size could affect the powder thermal conductivity and their absorptivity as well.in[12,13]. The melt pool width shows acceptable agreement with the experimental observations,in Fig. 8. Likethe predicted melt pool depth the penetration depthis underestimated comparedresults,increases when power1716(4)Figure 7: Temperature distribution for P=150 Watt, and V=1200 mm/s.a) P=100 Wattb) P=150 Wattc) P=195 WattFigure 8: Thermal results of single bead analysis for different process conditions1717The model was extended to a larger case where a bidirectional and chessboard scanning strategies were adopted. A block size of 6mm × 6mm × 1mm was considered with hatch spacing of 0.1 mm. The scanning pattern for the bidirectional was shown in Fig. 4(a). The temperature distribution and the melt pool size for the bidirectional and the chessboard scanning strategies are shown in Fig. 9 and Fig.

4(a). The temperature distribution and the melt pool size for the bidirectional and the chessboard scanning strategies are shown in Fig. 9 and Fig. 10, respectively. For both temperature distribution from the previous scanning vector. It was observed that the melt pool width and depth during the second scan are larger than the one in the first scan. The predicted residual stresses for the bidirectional scanning strategy are shown in Fig. 11. Within depth 100 µm, it was found that the principal component of residual stresses is 108±47 MPa. This result can find an agreement with the measured residual stresses for similar scanning strategy and process parameter in [3].

This result can find an agreement with the measured residual stresses for similar scanning strategy and process parameter in [3]. Although using coarser mesh for the structural analysis is computationally efficient, a fine mesh at high temperature gradient regions is needed to allow capturing that temperature variation.scanningstrategies,the melt poolsizeisinfluenced bytheFigure 9: Temperature distribution for bidirectional scanning strategy for a large block (6 mm × 6 mm × 1 mm) (P=200 W, V= 650 mm/s)Figure 10: Temperature distribution for chessboard scanning strategy for a large block (6 mm × 6 mm × 1 mm) and 2 mm chessboard block length (P=200 Watt, V=1000 mm/s)1718Figure 11: Induced residual stresses (principal component) after bidirectional scanning strategy4. Conclusion and Future WorkThis paper proposes a thermo-mechanical model for modeling the SLM process for scanning Ti-6Al-4V using strategies.

Conclusion and Future WorkThis paper proposes a thermo-mechanical model for modeling the SLM process for scanning Ti-6Al-4V using strategies. As exhibits reasonable accuracy and is also capable of handling large domain size with relatively high computational efficiency. The model predicts the melt pool width accurately but underestimates the melt pool depth given available experimental data. The reason could come from the model assumptions and simplifications. Hence, a systematic model validation study will be followed after this study. In addition, various uncertainties in model process parameters may greatly affect the thermal and mechanical properties of the printed products. Hence, these influences will be studied as well especially from the reliability and quality perspective.the comparedre-meshing totechniquethat experimentaladapts with different the modelavailableresults,5.

nite element simulation of direct deposition additive manufacturing for fiber reinforced thermoplastics Zhaogui Wang1, Zhenyu Fang1, Douglas E. Smith2 1Department of Mechanical Engineering, Dalian Maritime University, Dalian, China, 116000 2Department of Mechanical Engineering, Baylor University, Waco, TX, USA, 76798 AbstractPolymer deposition Additive Manufacturing (AM) (otherwise known as fused filament fabrication or fused deposition modeling) is extending to large format manufacturing due to the implementation of efficient screw-extruder-based material feeding systems [1]. Due to the low-cost and high capability in building intricate structures, the large scale AM systems have seen continuously applications in automotive, aerospace and naval industries for rapid prototyping of large-dimension parts and tooling [2-5].

Due to the viscoelastic nature of the thermoplastic-based feedstock, molten polymers are able to be deposited freely in 3D space without support material underneath. This feature facilitates the rapid freeform fabrication of grid cells that exhibit lightweight feature yet high structural strength and stiffness. For instance, Branch Technology Inc. (Chattanooga, TN, USA) mounted a single screw extruder onto a robotic arm and deposited Carbon Fiber filled ABS (CF-ABS) polymers directly in open space forming the lattice structures as appearing in Figure 1 [6]. During the deposition process, the extruded bead rapidly transits from viscoelastic state to elastic state and thus provide sufficient stiffness for the cellular structure to be built upwards. Carbon fiber filled thermoplastic composites provide additional stiffness and low thermal expansion that smooth the process of direct deposition additive manufacturing.

Carbon fiber filled thermoplastic composites provide additional stiffness and low thermal expansion that smooth the process of direct deposition additive manufacturing. In particular, the discontinuous fibers re-orient as the material flow being extruded and stretched in the process of direct deposition and thus the resulting fiber orientation in the beads affects the material properties of the manufactured grid cells. Nixon, et al. evaluated the internal geometric effects of the extrusion nozzle on the fiber orientation alignment using Moldflow software (MoldflowNon-planar direct deposition additive manufacturing of polymer composites has recently seen increased attention from industries, where the molten plastics beads are extruded through a heated nozzle directly into three-dimensional space to form lightweight truss-like structures.

This promising method shortens the manufacturing time and fabricates structures without stair-effect appearance that occurs in traditional layered additive manufacturing processes. This paper investigates the flow dynamics of fiber reinforced thermoplastics melt during a direct deposition process via a 2D planar deposition flow model. A quasi steady state for the deposition flow is assumed, where the flow-induced fiber orientation is evaluated through the Folgar-Tucker isotropic rotary diffusion model with the orthotropic-fitted closure approximation. The direct deposition process of a cubic lattice is simulated using finite element suite ABAQUS, where the orientation-homogenized material properties of a 13% CF-ABS are employed. Computed results indicate printing sequence affects the stress accumulation of the printed part notably.

Computed results indicate printing sequence affects the stress accumulation of the printed part notably. IntroductionSolid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference1366Governing equations for numerical modeling of fiber reinforced composite material flow for direct deposition AM are introduced in this section, including the identifications of the material flow fields as well as the fiber orientation state within the direct deposition flow suspension.Corporation, Framingham, MA, USA), where a convergent nozzle geometry was found to yield higher fiber alignment than a divergent designed nozzle [7]. Heller, et al. employed an axisymmetric 2D flow model via COMSOL (COMSOL Inc., Burlington, MA, USA) simulating the nozzle flow with a short section of free extrudate included, and found that the die swell at nozzle exit reduced the principal fiber orientation by ~20% [8].

Extended works based on Heller’s model explored the non-Newtonian fluid rheology effects [9] and the screw swirling motion [10] on the resulting fiber orientation of an extrudate composite. Studies reviewed above are done following a weakly coupled formulation between the flow kinematics and fiber orientation evolution, such that the fibers’ presence in the flow is ignored during the computation of the flow. A few studies performed fully coupled flow/orientation simulation for polymer deposition additive manufacturing, where depicted the mutually dependent interactions between the flow kinematics and fiber orientation kinetics (e.g., [11-13], while the fully coupled studies are mostly done for planar deposition flow. To reduce the complexity of the simulation, we start the flow modeling with the weakly coupled flow/orientation simulation, where the flow kinematics of a 2D Newtonian direction deposition (i.e., non-planar deposition) flow are solved firstly in ANSYS-Polyflow.

Based on the solved flow kinematics, the fiber orientation state of the flow domain is then evaluated using the Advani-Tucker orientation tensor approach through the Folgar-Tucker fiber orientation evaluation equation with isotropic rotary diffusion and orthotropic fitted closure. Employing the solved fiber orientation tensors at flow end, the material properties of a deposited bead are estimated through the orientation homogenization method, including the elastic constants, coefficients of thermal expansion and thermal conductivity. Finally, the transient direct deposition process of a simple cubic lattice is simulated via ABAQUS with computed properties employed. Figure 1. Grid lattice structures produced via direct deposition AM.

Jeffery’s Equation [16] first expressed the motion of a single rigid ellipsoidal particle. Folgar and Tucker [17] extended Jeffery’s theory to analyze the interactions between fibers in a non-dilute fiber suspension. Further, Advani and Tucker [18] defined the fiber orientation tensor approach to quantify the fiber alignment state for concentrated suspension systems, which requires fewer independent variables than that of the Folgar-Tucker model[17]. The Advani-Tucker equation can be written as [18] 1368Heller, et al. [8] (see also [9,10]) reported that the internal geometry of the nozzle has a notable impact on the fiber orientation of the extruded composite, since the flow domain geometry defined by the interior of the nozzle determines the melt flow velocity gradient fields. To this end, we consider the nozzle melt flow simplified as a 2D planar model, including the flow inside the nozzle and a short strand of free extrudate, as appearing in Figure 2.

Direct Deposition Flow Modelling1369Figure 2. Boundaries of the flow domain of interest. The transient velocity kinematics along flow streamlines are plotted through Figures 3-6 with a time increment of 0.05 second, in which the velocity variations during the deposition process can be viewed. It is seen that the flow end exhibits a similar velocity contour as long as the direct material deposition speed equals to the nozzle vertically-moving speed. Therefore, we assume a quasi-steady state of the fiber orientation could be achieved at the end of the flow. Further, the second order fiber orientation tensors are evaluated along the streamline data of simulation at t = 0.2 second (cf. Figure 6), where the quasi-steady state of fiber orientation is assumed to be achieved at the flow end of the domain (i.e., contact surface between flow end and substrate).

Figure 6), where the quasi-steady state of fiber orientation is assumed to be achieved at the flow end of the domain (i.e., contact surface between flow end and substrate). In the fiber orientation computation, the fiber aspect ratio is assumed as 15, in an averaged sense, which is a safe estimation based on prior studies on similar composite materials (e.g., [20,21]). The  is set as 0.01, as did in a similar study [10]. The orientation tensors diagonal components are given in Figures 7-9 (cf. flow streamlines locations are shown in Figure 10). It is seen that the fibers re-orientate intensely as being extruded out of the nozzle (i.e., at locations where horizontal coordinates of Figures 7 to 9 are around zero) and quickly recovers to a stable status as directly deposited onto the substrate (i.e., horizontal coordinates proceed to -0.02), which implies that the fibers could reach a steady state at the flow end.

The fiber orientation tensors diagonal components at flow end across the radial direction of the nozzle is given in Figure 11, which is then employed in the orientation homogenization approach [18] to predict the material properties of solidified deposited composites. The carbon fiber and ABS matrix properties are given in Table 1, and the computed properties of 13 wt.% CF-ABS (i.e., volume fraction is about 8.4%) are given in Table 2 sequentially. The computed properties appearing in Table 2 will be employed in the macro-level direct deposition AM process simulation shown in next section.1370Figure 3. Direct deposition flow at t= 0.05 second. Figure 4. Direct deposition flow at t= 0.10 second. Figure 5. Direct deposition flow at t= 0.15 second. Figure 6. Direct deposition flow at t= 0.20 second.1371Figure 7. Fiber orientation tensor component A11 along flow direction. Figure 8. Fiber orientation tensor component A22 along flow direction. Figure 9.

The printing process is divided into an element-by-element activation process, where the stem time for activating each element is 0.1 second. The boundary conditions are set as appearing in Figure 12, where the substrate and ambient temperatures are 40 ℃ and the heat convection coefficient is 82 /( ), adopting from [22]. Specifically, it is important to note that the deposited material exhibits a high degree of material anisotropy, as appearing in Table 2. In direct deposition AM, the principal direction of the beads’ anisotropy (maximum magnitude of the properties among the Cartesian coordinates directions) is determined by the direction of deposition. To this end, we employ different material orientation local coordinates based on the material loading direction, as seen in Figure 12 (i.e., coordinates 11-22-33).

To this end, we employ different material orientation local coordinates based on the material loading direction, as seen in Figure 12 (i.e., coordinates 11-22-33). Besides, the dimensions of a unit deposition element is 3 mm in thickness, 8 mm in width, 8 mm in length, representing a 3 mm nozzle depositing beads with a moving speed at 80 mm/s. The dimension of the printed cubic cell is 40mm*40mm*24mm. Figure 12. Finite element simulation setup for a simple cubic lattice direct deposition process.1373The temperature field of the printed cell is shown in Figures 13 and 14, where the conditions in as-printed state and cooled down state are given, respectively. Herein, the cooling time is set as 300 second as large-dimension extruded beads require longer time to reach the ambient temperature.

Herein, the cooling time is set as 300 second as large-dimension extruded beads require longer time to reach the ambient temperature. The fixed displacements boundary conditions applied on the bottom nodes are not removed during the cooling post-processing, indicating that the cubic cell is cooled without being taken away from the printing substrate. The substrate and convection ambient temperatures are set as 40 ℃ assuming the printing is performed inside a temperature-controlled furnace. In Figure 13, it can be seen that the temperature decreases slowly during the printing process due in part to the large-dimension of a single extrusion unit. This brings up an issue where the vertically deposited polymers remaining in relatively high temperatures are still in semi-molten status and may not be able to provide enough stiffness to support the upper deposited beads.

Herein, we note that a further study on printing/cooling balance is needed to ensure these vertically deposited beads are reduced below Tg temperature before the next planar beads are deposited. Nevertheless, we expect that the vertically deposited beads could support the upper planar deposited beads as the overall print size is still limited in small scale. After cooled down by 300 seconds, the overall cubic cell roughly reaches the ambient temperature, as seen in Figure 14. The bottom beads exhibit lower temperature than the upper beads (i.e., no support underneath) as the heat conduction between the ABS composite and material substrate is more efficient in heat exchanging as compare to the heat convection between the extruded beads and the air flow. Additionally, the vertically deposited beads are in relatively higher temperature among the entire printed domain.

Additionally, the vertically deposited beads are in relatively higher temperature among the entire printed domain. Furthermore, we plot the Mises stress fields and the deformation fields in Figures 15 and 16 and Figures 17 and 18, respectively. And the results are also given in as-printed and cooled-down conditions. During the printing process (cf. Figure 15), the maximum stress occurs at the bottom beads that contacting the substrate, in which the beads are restricted by the substrate and thus thermal residual stress accumulates. This is similar to a traditional layered manufacturing process. Nevertheless, the maximum stresses are seen in the vertically deposited beads in the cooled down condition (cf. Figure 16), and the magnitude of the maximum stress increases two orders as compared to that in as-printed condition.

Figure 16), and the magnitude of the maximum stress increases two orders as compared to that in as-printed condition. In addition, from the contours shown in Figures 17 and 18, it is seen that the maximum deformation occurs in the upper beads of the cubic cell, and increases one order of magnitude after being cooled down. The unique stress concentration and maximum deformation implies that the optimization of thermal post-processing is necessary to reduce the potential defects in the direct deposited lattice structures, which is expected to be explored by our presented simulation approach in other in-depth studies. Figure 13. Temperature field of the as-printed cell. Figure 14. Temperature field of the cooled-down cell.1374Figure 15. Mises stress field of the as-printed cell. Figure 16. Mises stress field of the cooled-down cell. Figure 17. Deformation field of the as-printed cell. Figure 18. Deformation field of the cooled-down cell.

Figure 16. Mises stress field of the cooled-down cell. Figure 17. Deformation field of the as-printed cell. Figure 18. Deformation field of the cooled-down cell. ConclusionA two-step numerical study is performed to simulate the direct deposition AM process. The weakly coupled flow/orientation analysis is adopted for a direction deposition AM nozzle flow. The transient extrusion deposition flow kinematics are solved by ANSYS Polyflow, where a quasi-steady state is assumed to be achieved at the end of the transient simulation. The Advani-Tucker approach is employed to evaluate the second order orientation tensors along streamlines of the modelled flow. Evaluated orientation tensors at the flow end are adopted to the orientation-homogenization-based properties evaluation to estimate the material properties of the 13% CF-ABS material extrudates.

Evaluated orientation tensors at the flow end are adopted to the orientation-homogenization-based properties evaluation to estimate the material properties of the 13% CF-ABS material extrudates. Direct deposition AM of a simple cubic cell is simulated by using the ABAQUS suite, where the transient thermal-mechanical history of the printed part is computed. The results indicated that the direct deposition is a promising AM approach for freeform fabrication of lightweight lattice grid-cells, while the printing/cooling timing balance is needed to ensure enough stiffness for upper deposited beads can be provided from vertically deposited beads. In addition, larger value of residual stresses is seen in vertically deposited beads as compared to planar deposited beads. To this end, we suggest to optimized post-cooling processing to reduce the thermal stress and deformation occurrences.1375[1] Love, Lonnie J., et al.

Ph.D. Thesis, Univer-sity of Illinois at Urbana-Champaign, Champaign, IL, USA, 1991. [20] Ning F, Cong W, Qiu J, et al. Additive manufacturing of carbon fiber reinforced thermoplastic composites using fused deposition modeling. Composites Part B: Engineering, 2015, 80: 369-378.AcknowledgementsThe authors would like to thank Dalian Maritime University and Baylor University for the support on this project. This work was supported by the Fundamental Research Funds for the Central Universities of China (No. 3132021337). References1376[21] Wang Z, Smith D E, Jack D A. A statistical homogenization approach for incorporating fiber aspect ratio distribution in large area polymer composite deposition additive manufacturing property predictions. Additive Manufacturing, 2021, 43: 102006. [22] Zhang Y, Chou Y K. Three-dimensional finite element analysis simulations of the fused deposition modelling process.

NITE ELEMENT STRUCTURAL ANALYSIS OF PROSTHESIS SOCKETS FOR BELOW-THE-KNEE AMPUTEES MANUFACTURED USING SLSMario Faustini*‡, Mahendra Lokhande*, Richard Crawford*, William Rogers†, Andrew Gitter† and Gordon Bosker†The University of Texas at Austin †The University of Texas Health Science Center in San Antonio ‡Supported by the National Council of Scientific and Technological Development (CNPq-Brazil)AbstractA very attractive application of Solid Freeform Fabrication is manufacture of prosthesis sockets for below-the-knee amputees. The custom geometric design required is very compatible with SFF techniques. The present work focuses on finite element analysis of sockets manufactured by Selective Laser Sintering using Duraform as the material. The objective is to ensure reliability of the sockets for their use by patients.

The objective is to ensure reliability of the sockets for their use by patients. This paper describes the finite element models developed for the sockets, as well as the derivation of realistic boundary conditions that may allow a simulation of the structure under regular workloads.1) IntroductionThe use of Selective Laser Sintering (SLS) in the manufacture of prosthesis sockets for below-the-knee amputees has been recognized as an appropriate potential application of Solid Freeform Fabrication [Rogers et al., 1991]. However, to ensure the reliability of these sockets and their safety for users requires a complete structural analysis of their mechanical behavior during operation. The focus of the current work is a finite element (FE) evaluation of the performance under standard conditions of one sort of socket, the single-wall socket.This paper is structured as follows: first, the single wall socket design is presented, followed by the development and validation of its FE model.

Next, the standard workloads during gait are defined as boundary conditions for the FE analysis. Finally, the results of the simulations of two sockets that were actually manufactured are exhibited.2) The single wall socketThe single wall socket, as seen in Figure 1, consists of an upper part, where the residual limb fits, and a lower region, which is a hollow cylinder where a pylon is attached with a prosthetic foot at its other end. The first step in the design and manufacture of the socket is fabrication of a mold from the patient’s residual limb, from which a copy of the limb is obtained. This copy is digitized with a laser scanning system. From this data, a perfectly fitting model of the socket is designed using computer-aided geometric operations on the data mesh. The thickness of the socket wall is decreased around specific points to provide compliance in those areas.

The thickness of the socket wall is decreased around specific points to provide compliance in those areas. The compliant regions relieve pressure on sensitive spots of the patient’s limb, such as the fibular head and distal tibia regions, thereby increasing comfort. Thus, tessellated data file (STL) is generated from this model and transmitted to a Selective Laser Sintering workstation for381fabrication. The material used is Duraform, the physical properties of which are given in Table 1.Figure 1. The single wall socketTable 1.

This level of detail would potentially result in unnecessary and time-consuming data preparation and analysis. Thus, the first step to generate the FE model is to apply a reliable mesh decimation algorithm to the STL file. We use Amira1 to perform the decimation operation. As seen in Figure 2, the decimated version has 2000 faces, but the overall geometric details and features are well preserved (especially the thicknesses of the wall, which is fundamental for the structural analysis).Figure 2. A section of the original STL model (48000+ faces) and the decimated version (2000 faces).Once decimated, the STL file is loaded into Rhinoceros2, where the triangular faces of the model are converted to parametric surfaces (NURBS) and exported as an IGES file. The IGES format is readable by SDRC I-DEAS3 for FE analysis.

The IGES format is readable by SDRC I-DEAS3 for FE analysis. In I-DEAS, the volume is meshed with solid elements (parabolic tetrahedra), assigned material properties and appropriate boundary conditions, and analyzed.4) FE Model ValidationIn order to validate the results of the FE structural analysis, an actual socket was tested to determine the vertical load that causes mechanical failure. The test is illustrated schematically in Figure 3.1 Amira is 3D Visualization package developed by Indeed-Visual Concepts GmbH, Berlin, Germany. 2 Rhinoceros is a 3D modeling package offered by McNeel & Associates, Seattle, WA. 3 I-DEAS is a CAD/CAE/CAM package provided by SDRC, Milford, OH.383Figure 3. Experimental loading test to determine the maximum vertical load a socket supports.The results were computed, as the load was gradually increased, in a force–strain graphic,shown in Figure 4. The figure indicates that the socket fails at a test load of 10 kN.Figure 4.

The figure indicates that the socket fails at a test load of 10 kN.Figure 4. Results of the experimental loading test.Similarly, a FE model equivalent to the physical model in Figure 3 was analyzed. Using the properties shown in Table 1, it was determined through repeated simulations that the socket reaches the ultimate tensile strength for Duraform (44 MPa) under a load of 6 kN. The simulation results for a 6 kN vertical load are shown in Figure 5.384Figure 5. Results for the simulation of the loading test using FE analysis.The difference between the ultimate load for the empirical test and the FE analysis may be explained by the fact that the socket used on the first one was infiltrated with cyanoacrylate, which increases the tensile strength of the part.

Qualitatively, however, the location where the actual socket cracked matched exactly the region of largest stress predicted by the FE analysis.5) Finite Element Analysis of GaitIn order to simulate the structural behavior of the FE model of the socket during gait, it is necessary to evaluate the loads applied by the residual limb. Figure 6, from Noguchi et al. [2000], shows the components of the ground reaction force of the foot in one stride (with z being the direction of the weight and x the direction of progress) in terms of percentage of body weight. Note that during the double supporting period (between heel contact of one foot and toe contact of the other (cid:190) when both feet are in contact with the ground), the forces have their greatest values, especially in positions A and B. Therefore, those two positions were analyzed in a worst case scenario, with all forces applied exclusively to the prosthetic foot.385Figure 6.

Therefore, those two positions were analyzed in a worst case scenario, with all forces applied exclusively to the prosthetic foot.385Figure 6. Components of ground reaction force of feet during gait [Noguchi et al., 2000].The boundary conditions for the FE model were calculated using the values for the forces in both positions A and B, as shown in Figure 7. In the figure, Fr is the horizontal resultant force due to Fx and Fy and Mr is the moment needed to balance the forces transferred from the ground to the middle of the socket.Figure 7. Boundary conditions used in the gait FE simulations.3866) SimulationsTwo sockets were simulated for both positions A and B (Figure 6) using the FE analysisprocedure described in Section 3. The patient was assumed to have a body mass of 75 kg.The first socket analyzed had a notch in the junction between the base cylinder and the lower thinner part in the front of the socket, as seen in Figure 8.

This socket was actually fabricated and broke while the patient was wearing it. The simulation results for this socket in position A are presented in Figure 9 (stresses are given in Pa). The simulation results for the same socket in position B are presented in Figures 10 (stresses in Pa) and 11 (displacement in meters).Figure 8. First socket simulated.Figure 9. Predicted stresses for first socket in position A.387Figure 10. Predicted stresses for first socket in position A.Figure 11. Predicted displacements for first socket in position B.388The second socket analyzed was designed with software with improved smoothing features, and is shown in Figure 12. This socket was also fabricated and has been used by the patient without failure. The simulation results for this second socket in position A are shown in Figure 13 (stresses in Pa).

This socket was also fabricated and has been used by the patient without failure. The simulation results for this second socket in position A are shown in Figure 13 (stresses in Pa). The simulation results for the same socket in position B are presented in Figures 14 (stresses in Pa) and 15 (displacement in meters).Figure 12. Second socket simulated.Figure 13. Predicted stresses for second socket in position A.389Figure 14. Predicted stresses for second socket in position B – bottom-left-front view.Figure 15. Predicted displacements for second socket in position B.390It should be pointed out that for both sockets the greatest stresses were computed for position B, even though the load values for position A were slightly larger (see Figure 6). Also, the analysis predicted, for position B, greater maximum stress for the first socket (15.5 MPa) than the second socket (14.6 MPa).

Also, the analysis predicted, for position B, greater maximum stress for the first socket (15.5 MPa) than the second socket (14.6 MPa). However, this still does not explain the failure, since this maximum stress is considerably less than the yield strength for Duraform (37 MPa – Table 1). Additionally, the maximum displacement for each socket was less than 2 mm, which preserves the biomechanics of the gait [Stephens, 1999].7) ConclusionThis paper presents a procedure for finite element based structural analysis for a below- the-knee prosthetic socket. Verification of the FE model and the simulation of two actual sockets are discussed. While the results agree qualitatively, the actual numerical results differ significantly. To address this disagreement, a more precise FE simulation algorithm will be tested (p-adaptive method). Also, more precise material properties for Duraform must be obtained.

To address this disagreement, a more precise FE simulation algorithm will be tested (p-adaptive method). Also, more precise material properties for Duraform must be obtained. Both Thompson [1995, 1997] and Watson [1999] illustrated the orthotropic behavior of Duraform, which depends highly on the direction of the parallel layers of the material in the structure (defined during manufacture) with respect to the direction of the stresses applied. Finally, both impact and fatigue studies should be addressed in order to provide a more accurate structural analysis of the socket, as these modes of failure are likely to occur in real sockets.8) References1. Rogers, W. E., Crawford, R. H., Beaman, J. J., and Walsh, N. E., 1991, “Fabrication of Prosthetic Sockets by Selective Laser Sintering,” 1991 Solid Freeform Fabrication Symposium Proceedings, Marcus, H. L., Beaman, J. J., Barlow, J. W., Bourell, D. L., and Crawford, R. H., eds., Austin, TX, August 12-14, 1991, pp. 158-163.2.

nite Element Thermal Analysis of Three Dimensionally Printed (3DP™) Metal Matrix CompositesS. Johnston1,2 and R. Anderson2*1 Dept. of Mechanical Engineering, University of Washington, Seattle, WA 98195 2 Concurrent Technologies Corporation, Bremerton, WA 98312 *Author to whom correspondence should be addressed.ABSTRACTThree Dimensional Printing, 3DP™, is under development for the production of metal composites through solid freeform fabrication of a skeleton, followed by sintering and then infiltration of a second, lower melting metal. To maintain dimensionality and allow for infiltration, green component sintering produces only neck growth between powder particles, which differs substantially from conventional powder metallurgy sintering practices. The objective of this study is to investigate finite element thermal-mechanical modeling in order to optimize the sinter profile and predict the final part dimension.

The objective of this study is to investigate finite element thermal-mechanical modeling in order to optimize the sinter profile and predict the final part dimension. The model will relate micro-scale thermal effects to the macro-scale by analyzing the composite medium as a meshed layered element comprising micro-scale elements. The micro-scale elements can be based on bulk transport mechanisms related to neck growth. The analytical model will eventually be used to define the thermal processing parameters and to predict optimal geometry for full density near net shape sintering and infiltration of 3DP™ manufactured components.INTRODUCTIONThree Dimensional Printing, 3DP™,1 is a rapid prototyping and manufacturing process that creates complex three-dimensional parts directly from a computer-aided design (CAD) model.

Developed at the Massachusetts Institute of Technology, Three Dimensional Printing is one of several solid freeform technologies for the production of prototypes, functional parts, and tooling. Some of the benefits of the unique technology of 3DP™ include the economic fabrication of small quantity parts and one-offs, the use of alternative materials, and the redesign of parts and assemblies to enhance performance and durability.The 3DP™ powder metallurgical process involves spreading a thin layer of metal powder over the surface of a part-build bed. Using ink-jet printing technology, a liquid binder material is deposited to selectively join the powder particles where the part is to be formed. The powder bed and part are lowered and a new layer of powder is spread and selectively joined. This layer-by- layer printing process repeats until the part is completed.

The powder bed and part are lowered and a new layer of powder is spread and selectively joined. This layer-by- layer printing process repeats until the part is completed. After thermal processing for debinding and sintering, the fabricated preform, or “skeleton,” undergoes infiltration to form a metal matrix composite component that is capable of serving not only as a prototype but, depending on the material system and application requirements, as a fully functional part.Understanding the thermal processing mechanisms during sintering is essential todimensional and surface quality control of 3DP™ components. In contrast to machined, cast, or forged components, 3DP™ powder metallurgical parts require three specific heat treatments before the component is ready to be utilized: (1) curing of the binder following printing, (2) sintering of the green components and (3) infiltration. Each of the three thermal processes cause432dimensional variances.

Each of the three thermal processes cause432dimensional variances. The objective of this study is to investigate finite element thermal- mechanical modeling in order to optimize the sinter profile and predict the final part dimension. The analytical model will eventually be used to define the sintering parameters and to determine, in the design stage, if a part geometry can be satisfactorily sintered. This paper discusses the sintering portion of the thermal aspects of the 3DP™ process.Overview of SinteringThe microstructural evolution that occurs during sintering involves three distinct phases priorto achieving full component densification. The initial stage involves neck growth between powder particles which results in a light cohesive bond. The intermediate stage involves pore rounding, the densification of the sinter compact, and grain growth.

The intermediate stage involves pore rounding, the densification of the sinter compact, and grain growth. The final stage involves the closing of pores and pore shrinkage, resulting in total densification.2The complexity of sintering has yet to be completely theoretically defined, thus sintering is still referred to as a phenomenological process.2,3 The difficulties with creating suitable sintering models involve lack of data at sintering temperature for materials, non-existent constitutive equations relative to sintering, and model verification (despite use of advanced mathematics). For example, isothermal models do not accurately model the sintering process because dilatometry experiments show that sintering occurs while ramping to the peak temperature. The level (macro- or micro-) at which the computer simulation places emphasis is also important.

The level (macro- or micro-) at which the computer simulation places emphasis is also important. Concentrating the simulation of the sintering process at the micro-scale versus the macro-scale depends upon the desired quantities to be measured. Ultimately, a combination of an atomistic model (i.e., micro-level) and a finite element model (i.e., macro-level) is held by many researchers to be a useful, accurate, simulation approach.4,5Initial Stage Sintering MechanismsInitial stage sintering produces neck growth through two different classes of sinteringmechanisms: surface transport and bulk transport. Since it is possible to calculate thermal strains from measured shrinkage values, a logical starting point for the model is to focus on the bulk sintering mechanisms that produce physical shrinkage. These mechanisms include plastic flow (PF), grain boundary diffusion (GB), and volume diffusion (VD).

These mechanisms include plastic flow (PF), grain boundary diffusion (GB), and volume diffusion (VD). Each mechanism is active during different stages of the sintering process so that the total shrinkage produced is the superposition of these three bulk mechanisms.2,6Plastic flow can be described as nonconservative dislocation motion. It is directly related tothe temperature profile and is active while the component’s temperature is approaching the sintering temperature, so PF is only present during initial stage sintering and ceases once the sintering temperature is achieved.2 The more severe ramp in temperature profile, the greater the rate in plastic flow.Grain boundary diffusion involves the realignment of grain boundaries due to crystalmisalignment. This mechanism depends primarily upon the grain size and concentration of grain boundaries per unit volume.

This mechanism depends primarily upon the grain size and concentration of grain boundaries per unit volume. The movement of atomic particles along the grain boundaries usually possesses an activation energy less than that of volume diffusion, and therefore is not as influential in shrinkage as volume diffusion.433Volume diffusion involves the movement of particles through lattice vacancies of acrystalline structure. The diffusion rate is dependent upon three parameters associated with the powder structure: temperature, composition, and pressure (or curvature). The evolution of neck growth from volume diffusion is essentially the redistribution of atoms within vacancy positions that leads to mass distribution between the particles.

The evolution of neck growth from volume diffusion is essentially the redistribution of atoms within vacancy positions that leads to mass distribution between the particles. Temperature is the dominating factor controlling the concentration of vacancies; hence, the rate at which material mass may reposition is related to temperature.The necks that form between particles during sintering produce the effective straincontribution for each of these bulk mechanisms. The neck growth and shrinkage between two equally sized spheres is represented in Figure 1 (see also Equations (1) and (2), and Table 1).Figure 1.

The neck growth and shrinkage between two equally sized spheres is represented in Figure 1 (see also Equations (1) and (2), and Table 1).Figure 1. Schematic representation of a neck growth model for two equally sized spheres.The effective strain (εi) for each mechanism follows a kinetic law as shown in Equation (1):( ε i)n2/= ∆ L  L   n2/=2Bt n Dmwhere t is time, D is particle diameter, L is the original length between the centers of the spherical particles, ∆L is the change in distance between centers, and the remaining parameters are defined in Table 1. The constant B in Equation (1) is dependant on the specific sintering mechanism being considered, and is exponentially dependent upon temperature by the relationship shown in Equation (2):= BBo exp −  Q kT  where Q is the activation energy, and the remaining variables are defined in Table 1.Table 1.

The layer is subdivided into smaller elements comprising a number of powder particles modeled as tetrahedral elements whose deformation is governed by Equation (1) and the information in Table 1. The sintering bonds produced between layers (i.e., z-print direction) would have less particle contact due to the compaction of the particles in this direction. This is an inherent trait of the printing process and has been observed experimentally.EXPERIMENTAL RESULTSRectangular sample blocks were printed in 316L stainless steel (316L SS) and 420 stainlesssteel (420 SS) powder with dimensions of 25x25x40 mm, where the 40 mm dimension was oriented in each of the three (i.e., x-, y-, and z-) print directions. The sample blocks were measured in the green state and sintered state using a non-contact video measuring device (see Experimental Procedure).From these measurements, the thermally induced sintering strains were determined as the normalized change in length using Equation (3).

Results are shown for 316L SS and 420 SS in Figures 3 and 4, respectively.=ε T∆L LStatistical analysis (t-test) of the measured strains indicated that the strain produced in eachof the three print directions is independent of each other for both the 316L SS and 420 SS samples. Thus, the microstructural development of the component is anisotropic.70.0110.0130.015)m m m m/0.017( n a r ti0.019SAverage Strain0.0210.023X-DirY-DirZ-Dir0.025Print DirectionFigure 3. Sintering strains produced in 316L SS rectangular sample blocks.436(3)0.0040.0050.0060.007)m m m m0.008/( n i a r t S0.0090.01Average Strain0.0110.0120.013X-DirY-DirZ-DirPrint DirectionFigure 4. Sintering strains produced in 420 SS rectangular sample blocks.The mean strain for each print direction is represented by the trendlines in Figures 3 and 4.

Sintering strains produced in 420 SS rectangular sample blocks.The mean strain for each print direction is represented by the trendlines in Figures 3 and 4. The mean strain observed for each print direction is unique, and the differences in strain for the three print directions can be attributed to two factors inherent to the 3DP™ process: binder droplet size and print direction itself. Specifically, deviations in the x-direction are attributed to the start and stop of binder distribution; deviations in the y-direction are attributed to the jet width of the print head, which produces slight overlap for adjacent print y bands; and deviations in the z-direction are attributed to a combination of powder/binder adjacent layer adhesion and layer thickness. Further analysis of the 316L SS and 420 SS data reveals that the strain in each print direction shows small standard deviations.

Further analysis of the 316L SS and 420 SS data reveals that the strain in each print direction shows small standard deviations. The normal distributions in Figures 5 and 6 indicate that although unique to print direction, the strain values are reproducible and consistent.250200n o i t u b i r t s i Dl a m r o NXdirection Ydirection Zdirection1501005000.040.02 0 Sintering Strain (m m/m m)0.020.04Figure 5. Normal distribution of sintering strains for 316L SS sample blocks.437400n o i t u b i r t s i Dl a m r o NXdirection Ydirection Zdirection30020010000.040.0200.020.04Sintering Strain (mm/mm)Figure 6. Normal distribution of sintering strains for 420 SS sample blocks.CONCLUSIONSA 3DP™ manufactured component is anisotropic; that is, it has the quality of variation of physical properties with the (x-, y-, or z-print) direction along which the property is measured.

This anisotropic nature is related to aspects of the three-dimensional layered physical construction of the component, the effects of which are retained through thermal processing. This study introduced initial steps toward finite element thermal analysis to predict the effective strain produced from sintering 3DP™ manufactured components. The model utilizes a meshed layer procedure and incorporates both the micro-scale aspects of sintering between particles as well as the macro-scale parameters of dimensionality and strain. Due to the anisotropic construction of 3DP™ components, determination of strain in each print direction is the initial step to defining the thermally induced stresses within the component. The strains are used to predict localized stresses within the component, which in turn provides a prediction of possible positions of fracture.

The strains are used to predict localized stresses within the component, which in turn provides a prediction of possible positions of fracture. Consideration of the layer-by-layer construction inherit to the 3DP™ manufacturing technique is essential in order to correctly simulate as many process traits as possible, providing for a useful, accurate modeling tool for the 3DP™ process.EXPERIMENTAL PROCEDUREComponents were manufactured on a ProMetal RTS-300, a solid freeform fabricationmachine capable of creating metal components as large as 12x12x10 in. Typical process speed is 10-20 in3/hr, depending on resolution. Components were thermally processed in a nonvacuum atmosphere furnace. Metal powders and binder were used as obtained from commercial vendor (Extrude Hone Corporation, Irwin, PA).Dimensional measurements were obtained using a Smart Scope non-contact video measuringdevice (Optical Gaging Products, Inc.).

The Smart Scope measuring method selects several438points along each edge of the sample at a magnification of 53x. A least squares linear fit is used for the edges of the samples. Dimensional measurements are calculated from the linear fit lines with an accuracy of ±40µm (0.00015 in).ACKNOWLEDGEMENTSFinancial support for this work is provided by the Office of Naval Research, Contract #N00014-C-00-0378. The authors are grateful to Bill Pedersen (University of Washington) for assistance with experimental data interpretation, and the Physics Department at University of Washington for equipment use. We also thank Angela Wilson for graphic design assistance, Dawn Buckley and Melinda Webel for editorial assistance.REFERENCES1. 3DP is a registered trademark of the Massachusetts Institute of Technology.2. R. M. German, Sintering Theory and Practice, Wiley & Sons, New York, 1996.3. E. A.

3DP is a registered trademark of the Massachusetts Institute of Technology.2. R. M. German, Sintering Theory and Practice, Wiley & Sons, New York, 1996.3. E. A. Olvesky, Theory of Sintering: from discrete to continuum, Material Science and Engineering: R, Reports, Vol. R23, Issue 2, June 1998, pp.41-100.4. R. M. German, personal communication and presentation, Critical Overview of Sintering Computer Simulations, at The 2002 World Congress on Powder Metallurgy & Particulate Materials, Orlando, FL, June 2002.5. E. A. Olvesky, personal communication and presentation, Sintering of Functionally Gradient Ceramic Freeforms, and presentation, Sintering of Multilayer Powder Composites: Distortion and Damage Control” at the 2002 World Congress on Powder Metallurgy & Particulate Materials, Orlando, FL, June 2002.6. Powder Metal Technologies and Applications, ASM Handbook, Volume 7, 1998 pp.448-452.7. A. P. Boresi, R. J. Schmidt, and O. M.

RST STEPS TOWARDS COLLABORATIVELY EDITED DESIGN FOR ADDITIVE MANUFACTURING KNOWLEDGEE. L. Doubrovski, J. C. Verlinden, I. HorvathDepartment of Industrial Design Engineering, Delft University of Technology, Delft, the NetherlandsAbstractDespite the broad coverage concerning the technological challenges, little research has been performed on the methods that enable designers to deal with Additive Manufacturing. At present, the challenge is to generate Design for Additive Manufacturing knowledge which goes beyond traditional solutions and to ensure that this knowledge is complete, correct and up to date. This paper reports on the employment of a wiki environment to support open-ended knowledge management. We applied this solution in an undergraduate prototyping course focused on exploring visual properties using AM structures. The results of the 32 students encompass unexpected designs while the knowledge on the wiki encompassed i) AM processes, ii) procedures, iii) artifacts.

The results of the 32 students encompass unexpected designs while the knowledge on the wiki encompassed i) AM processes, ii) procedures, iii) artifacts. This forces us to rethink what should constitute DfAM.IntroductionAdditive Manufacturing (AM) is a maturing collection of production technologies also known as three- dimensional (3D) printing. Compared to conventional production, such as machining and injection molding, AM offers unprecedented possibilities in shape complexity and custom geometry. As the application of AM for the production of end-products is increasing, AM is having a growing influence on the product development cycle; the way products are designed, produced, distributed, consumed and recycled.While AM is rapidly developing, industrial designers are facing several challenges in applying this manufacturing technology. It has been stated that the advantages of using AM are not yet understood by designers [1].

It has been stated that the advantages of using AM are not yet understood by designers [1]. We argue that the issue is more fundamental: the current technological opportunities of AM are not yet fully known and, in addition, these possibilities are developing quickly, with new or improved processes and materials being developed on a regular basis. Within the AM research community, it is accepted that the development of Design for Additive Manufacturing (DfAM) methods is needed in order to allow designers to benefit from this manufacturing technology to the full extent [10]. However, despite the broad coverage of the technological challenges of AM, very little research has been performed on the methods that enable designers to deal with this game changer. Furthermore, among the DfAM approaches that have been proposed so far, very few have focused on the creative phases of the design process.

Furthermore, among the DfAM approaches that have been proposed so far, very few have focused on the creative phases of the design process. In the context of this paper, we consider conceptual design as the most creative phase of the design process. Figure 1 shows the environment of conceptual design in the broader scope of the industrial design process, as illustrated by Horvath [2].Figure 1: Environment of conceptual design891In developing a DfAM for the creative phases of industrial design we are facing several challenges. The rapid development of both the AM technology and the supporting software poses a risk of any DfAM method becoming obsolete within a short period of time. Furthermore, the possibilities of the currently existing technologies are still not fully discovered and documented.

Furthermore, the possibilities of the currently existing technologies are still not fully discovered and documented. This makes us pose the following question: How can we generate DfAM knowledge that covers techniques and methods which go beyond traditional solutions? Although different DfAM approaches are being proposed in literature, none of the approaches is based on a collaboratively edited platform. Addressing the issues described above, we aim to examine how designers who are exploring AM structures with new properties can provide valuable contributions to the knowledge on the possibilities of AM and whether a collaboratively edited platform can support the documentation and use of the new methods, procedures, and designs.

Demonstrating the feasibility of this approach, we present our results of an initial experiment among Industrial Design Engineering students using a wiki environment for collaborative editing of DfAM knowledge.Emerging approaches to DfAMOver the past years, as the application of additive technologies to production of end products has been increasing, several authors have identified the lack of design methodologies and design knowledge for AM, coining the term Design for Additive Manufacturing. Initially, most efforts to develop tools for designers included the development of support to select a suitable AM technology. Approaches ranging from knowledge- based systems [3] to neural networks and fuzzy logic [4] have been proposed to achieve this. Other methods that have been proposed support the AM production of components by assisting decision making such as choosing build orientation [5].

Other methods that have been proposed support the AM production of components by assisting decision making such as choosing build orientation [5]. Furthermore a reasonable amount of literature can be found on the development of processes and algorithms to optimize structures for AM in order to achieve specific properties [6]. Commercially available software systems are limited to data handling, 3D design checking and generation of cellular structures within a volume. Focusing on design rules, rather than design methodologies, the thesis of Thomas [7] describes the development of design rules for selective laser melting. His work describes the evaluation of the geometric limitations of one AM system (SLM) through quantitative cyclic experiment.

His work describes the evaluation of the geometric limitations of one AM system (SLM) through quantitative cyclic experiment. Based on the findings design rules were created which were then evaluated with design professionals.While the abovementioned approaches are intended as design support, none of them truly assist a designer during the creative phases of product development. Maidin & Campbell recognized this gap and have described one of the most extensive efforts to solve this problem [8]. Therefore, the authors of this paper aimed at developing a knowledge-based support tool for designers for the creation of products for AM. A database of AM design features is used as a knowledge repository. A collection of such AM design features has been derived from examples from literature. In order to make this knowledge accessible to designers, these design features have been organized into a taxonomy with four top-level taxons being the key reasons for using AM.

In order to make this knowledge accessible to designers, these design features have been organized into a taxonomy with four top-level taxons being the key reasons for using AM. This database has been tested with design students and professional designers, showing that it can be helpful and inspirational during the conceptual phase of a design process.A New ApproachThe database described by Maidin et al. [1], which is at the core of their proposed method, depends on a centralized identification of design features from the designer’s tacit knowledge and a centralized indexing of these design features into the database. However, the rapid global development of AM could make a centralized approach unmanageable. On one hand we see daily innovations in AM technology, such as materials, accuracy, surface finish, speed. On the other hand, we see a growing amount of new designs and smart applications of this AM technology.

On the other hand, we see a growing amount of new designs and smart applications of this AM technology. Therefore such a centralized scenario will pose challenges to the correctness, completeness and extensiveness of DfAM knowledge.We propose a complementary method based on collaboratively edited content. Current online examples such as Wikipedia and Instructables [9] illustrate that, if provided with adequate tools, international communities can build and maintain a vast and up-to date database of knowledge. A vast amount of insight is892being created by designers when designing and manufacturing products or parts with AM. Therefore, the worldwide community of designers who design for AM are potentially a fruitful basis for a collaboratively edited DfAM knowledge base. Although collaboratively edited knowledge has been in existence for over a decade, it has not yet been applied in the context of Design for Additive Manufacturing knowledge.

Although collaboratively edited knowledge has been in existence for over a decade, it has not yet been applied in the context of Design for Additive Manufacturing knowledge. We intend to explore if the documentation and usage of their knowledge can be effectively facilitated by providing designers with a wiki environment. As an initial test of our approach, a pilot study has been executed in the context of an educational module. We aimed to answer the following questions: How do designers interact with the wiki in terms of documentation and use?  What knowledge is generated on the wiki?Is a wiki a suitable environment for documenting and using DfAM knowledge?Pilot StudyThe pilot study has been performed during a 4 week course of a minor on Advanced Prototyping organized by the school of Industrial Design Engineering of the Delft University of Technology [10]. In total, 32 third year undergraduate students participated in the course, divided in four groups of 8 students.

In total, 32 third year undergraduate students participated in the course, divided in four groups of 8 students. All students had good knowledge of AM and CAD modeling. The course was organized in a so called “carrousel” structure; each team worked on one of four selected topics for one week before switching to the next topic. This means that each topic was tackled by a new group each week, in four consecutive weeks. Each group of 8 students was divided into teams of two students. Two of the four topics were directly related to AM: visual properties of AM, and AM cellular structures. The two other topics were related to other aspects of prototyping: reverse engineering, and “objet trouvé”.This study was performed within the carousel topic “visual properties of AM”. The assignment was relatively open, as the students were asked to explore techniques to achieve new visual and optical properties of AM-produced objects. In an overview of the literature, Doubrovski et al.

In an overview of the literature, Doubrovski et al. [6] have shown that functional improvements, such as product weight, are being covered extensively in AM research, but visual properties remain unexplored. Since such properties are arguably just as important for industrial designers as other functional properties, this was an interesting new topic to explore in the context of the course. Initially, three sub-topics were chosen as a starting point for further exploration of visual properties, these being reflectance, optical fibers, and subsurface scattering. As described by Doubrovski et al. [11], there lies an opportunity to manipulate these properties in AM produced components, but the adequate knowledge and design tools do not exist.As a platform for collaboratively edited content, the students were provided a wiki environment to document their progress, experiments, and results.

The free service of WikiDot.com [12] was chosen, which is a free, simple, and relatively easy to use wiki. We have created an initial structure of the wiki which consisted of a main “home” page that links to 3 sub-pages which cover the 3 determined sub-topics: reflectance, optical fibers, and subsurface scattering. These pages initially contained a general description of the sub-topic and several envisioned challenges. Students from the first three teams were assigned to do explorative work on the topic and to document their developments and gained knowledge on the wiki. Students from the final team were asked to make a design of a product, applying the knowledge available on the wiki. It was allowed to edit the wiki at any time, both the content and structure.

Students from the final team were asked to make a design of a product, applying the knowledge available on the wiki. It was allowed to edit the wiki at any time, both the content and structure. The students were not required to hand in any other form of written report apart from the wiki.For the evaluation of our approach of using a wiki for collaboratively edited DfAM, after the completion of the course we have analyzed the constructed content on the wiki, the editing behavior of the participants, as well as the participant reflections on the use of the wiki. The wiki was inspected, considering both the content and the structure. The editing behavior was collected through the logs of user activity on the wiki.

The wiki was inspected, considering both the content and the structure. The editing behavior was collected through the logs of user activity on the wiki. Finally, a questionnaire was used to collect the participant’s reflections of the use of the wiki at the end of the course.893Constructed knowledgeTaking into account that the course was initiated when the wiki contained almost no information, and that the students were unfamiliar with the topic of visual properties of AM produced components, in four weeks, the four consecutive groups of students have designed and built remarkable results and the wiki has grown into a collection resources on the selected sub-topics. Below we discuss the emerged structure and content of the wiki.StructureAt the end of the four-week course, the top level of the wiki structure has remained largely identical, with the sub-topics being present on the main page. The largest developments have taken place within the pages of the sub-topics.

The largest developments have taken place within the pages of the sub-topics. Figure 2 illustrates this structure of the wiki.Home: VisualOptical Fibres Co-extrusieSubsurfacescatteringReflectanceBusiness cardsUltimakerGroup 2 •Milou tests •Simone tests •Mariet tests •Kiran tests •Foto testsRenee & Anne Lien testsNick, Elian & Piet testsGroep 3 – UltimakerGroep 1 Plaatje in 3D pixelsGroep 4 testsGroep 3Figure 2: Emerged structure of the wiki.As can be seen in the overview, some students have created new pages that are linked to the pages of the sub-topics while others have edited and expanded the page of the sub-topic itself. For example, the sub-topic of “subsurface scattering” has 5 child pages while the page of the sub-topic “optical fibers” has all of the content on the page itself and has no child-pages. Considering the above, we observe that the emerged structure is inconsistent and could be confusing.

Considering the above, we observe that the emerged structure is inconsistent and could be confusing. Furthermore, we see that the naming of the created pages is inconsistent and in some cases uninformative, having the name of the participants or the group number. Additionally we have found some pages that are linked to from an incorrect location, for example the description of the smartphone case with reflectance properties is linked from the sub-topic “subsurface scattering”. The only form of active moderation by the authors has been the initial setup of the structure, no moderation was done during or after the project. The results described here indicate that this was not sufficient to maintain a clear and structured wiki.Knowledge levelsExamining the wiki after the completion of the course provides us an insight in the content that has been generated by the participants. We have observed that the wiki has accumulated DfAM knowledge on different levels.

We have observed that the wiki has accumulated DfAM knowledge on different levels. Most of the knowledge on the wiki concerns concrete knowledge about specific processes, materials and applications. For example, descriptions are provided on how specific components have been modeled using the Grasshopper plugin for the Rhinoceros 3D software package, including detailed instructions on how to reproduce the built geometries (Figure 3). These descriptions have been used and further elaborated by students from later groups.894Figure 3: Wiki content, describing settings for the Grasshopper plugin of Rhinoceros 3D.In addition to the processes descriptions of how specific visual properties were achieved, participants also provided descriptions of experiments on the wiki. These descriptions were valuable for successive groups who have applied the knowledge gained from these experiments.

These descriptions were valuable for successive groups who have applied the knowledge gained from these experiments. For example, members from the first group have tested different wall thicknesses of internal structures to see how light is passed through when using the VeroBlue material printed on the Objet Eden 260 3D printer, one of the test pieces is shown in Figure 4a. The test results, were used by members from the final group to design and 3D print a “beer glass cooler” that has internal features which result in a photograph with different shades when lit from behind (Figure 4b). Some participants have also added recommendations for further research on the visual phenomenon that they have explored.

Some participants have also added recommendations for further research on the visual phenomenon that they have explored. Also new ideas of applications of the visual properties have been described on the wiki pages.a b Figure 4: test piece (a) and beer glass cooler (b).Built models and designsAfter discussing the knowledge on the wiki, in this section we discuss the outcome of the experiments and design activities of the students. These designs have been documented on the wiki and therefore can function as a source of tacit knowledge.As already shown in figure 4a and 4b, several tests have been done whereby the internal geometry of a volume was designed to create different shades when the object is back-lit. Figure 5 shows a test piece in which a photograph can be seen.Figure 5: Embedded photograph in a 3D printed sample.895Another exploration into new visual properties was performed by manipulating the reflectivity of the 3D printed surface.

The workflow to achieve this is based on the processes used for the beer cooler, where the Grasshopper plugin was used to sample an image and use the grayscale values as a parameter for cylinder diameter. In this example, instead of the diameter, the angle of the top surface is altered according to a grayscale value. Figure 6 shows a smartphone case with a surface consisting of embossed cylinders. As each surface under a different angle reflects different amounts of light, the result shows an image of an apple in different shades.Figure 6: Smartphone case with reflective surface.Based on the same Grasshopper workflow, another approach to visualize images has been explored by modeling a grid of small tubes through a solid sheet. Figure 7 shows the result of a 3D printed backlit sheet (thickness 2mm), illustrating a portrait. This effect is achieved by altering the angle of each tube is dependent on the grayscale value of the representing pixel.

This effect is achieved by altering the angle of each tube is dependent on the grayscale value of the representing pixel. A full black pixel results in a 45 degree cylinder, while a white pixel results in a cylinder perpendicular to the plane surface.These examples illustrate how knowledge on the wiki, describing a workflow to achieve a specificresult, has been used to achieve new, completely different optical properties.Figure 7: Back-lit test piece showing a portrait896One of the other topics that was explored is the Additive Manufacturing of optical fibers. Although this topic was only covered during one carrousel cycle, some exciting results have been achieved. Figure 8 shows samples that have been produced by extruding PMMA optical fibers using a RepRap Fused Deposition Modeling machine.

Figure 8 shows samples that have been produced by extruding PMMA optical fibers using a RepRap Fused Deposition Modeling machine. The samples show that it is possible to produce objects with internal optical fibers allowing to direct light through a built volume.Figure 8: Optical Fibers Produced with Fused Deposition ModelingWiki useEditing behaviorBy inspecting the logged edits on the wiki, we were able to analyze the user activity. As each of the four groups worked for one full week on the topic, the amount of edits of each group is labeled according to the day of the weekly carrousel. Figure 9 shows the amount of edits per day per team. For each team, most of the editing activity has been recorded during the last day of the week, before finalizing the project.

Figure 9 shows the amount of edits per day per team. For each team, most of the editing activity has been recorded during the last day of the week, before finalizing the project. Two groups started earlier, with initial editing on the third day of the week.Figure 9: Page edits per team per day.The fact that most most of the editing and documenting was done at the end of each carrousel round could be explained by the logic that as knowledge and results of the teams grew they documented more of these results on the Wiki. However, this could also imply that students used the wiki to report their work in a similar manner that is common in design education. This is discussed in the next section. Unfortunately we did not have the means to track the amount of page views of the wiki, such data would provide a some more insight in the use of the wiki.QuestionnaireA questionnaire was handed out after the completion of the 4-week course to all students, 20 of whom responded.

The aim of this questionnaire was to capture the i) students’ reflection on the use of the wiki, ii) in which ways the wiki assisted the assignment, iii) how they think the wiki could be improved, and iv) how they envision future use of the emerged wiki. The questionnaire comprised of open questions, multiple choice questions and Likert scale questions. Because the first groups have worked with the wiki at the beginning of the897course, they were not familiar with the final state of the wiki. For this reason, before filling in the questionnaire, all students were asked to re-inspect the final wiki content. We discuss the main findings of the questionnaire below.Four statements were provided related to the use of the wiki. The respondents were asked to score their level of agreement with the statements. Each statement was to be scored on a 5-point Lickert scale. We consider the upper two Lickert levels a positive feedback.

Each statement was to be scored on a 5-point Lickert scale. We consider the upper two Lickert levels a positive feedback. The first statements were related to the use of the wiki during the project. Table 1 shows the percentages of respondents who have indicated to agree with the statements. We can see that the students reflect their use of the wiki mainly for documenting of their process, followed by the documentation of knowledge. The use of the wiki for the acquisition of knowledge is scored positive by 36% of the respondents.Wiki use for: Documentation of process Documentation of knowledge Acquisition of knowledge Collaboration with team membersReflected use (%) Envisioned use (%)82 72 36 2390 90 75 25Table 1: Reflected use of the wiki (n=20).After the respondents were requested to inspect the wiki as it was at the end of the course, similar statements as the ones mentioned above were asked, but now related to the envisioned future use of the wiki.

The responses show a similar ranking of agreement with the statements of envisioned use as the previous. We do see a significant increase of agreement with the statement that the wiki was used to acquire knowledge. Looking at the data we observe that during the project just 36% of the respondents have used the wiki to acquire knowledge for the project, while after looking at the more elaborated wiki, more than half of the respondents indicate to consider using this wiki for future projects. This increase in agreement is bigger for those, part of the first teams.Nearly 40% of the respondents have indicated that the wiki did not help them during the course. Unfortunately most of the students who have indicated this, have did not provide an explanation. Among the few reasons that have been mentioned are the arguments that the wiki was only used to document own work and not for retrieving information.

Among the few reasons that have been mentioned are the arguments that the wiki was only used to document own work and not for retrieving information. It is important to note that these responses were given mainly by students from the first two teams. In contrast, only 3 students from the last two groups have indicated that the wiki did not work for them. We can explain this by the fact that the use of the wiki was changing, as the amount of information culminated.Responses to the question concerning possible improvements of the wiki included the ease of use of the used wiki platform, such as embedding of images and video. Several students proposed adding a search function and a navigation menu.

Several students proposed adding a search function and a navigation menu. Although most groups have created a wiki page for their team, yet while reflecting on the final wiki content, most students indicated that the wiki would have a better structure if the pages would be structured according to content and not according to team.DiscussionEducational contextThe experiment covered a weekly assignment on structures and visual properties. This was carried out by 4 consecutive groups of students, all sharing the wiki to keep track of insights and findings. It is obvious that this situation does not resemble a regular design project, and the level of commitment is differing from an industrial context. Furthermore, the knowledge necessary for this assignment does not require integration with other types of engineering and design, normally one of the main challenges in product design.

Furthermore, the knowledge necessary for this assignment does not require integration with other types of engineering and design, normally one of the main challenges in product design. Taking this into account, we cannot guarantee that professional designers will have the same relationship with this approach and898further experiments with professional designers are needed. However, the knowledge gained during the experiment surpassed known solutions and allowed a hands-on learning experience. As a didactical exercise, it does include apprehending new insights as design does. Also, many of the results resemble products, such as the beer holder. With 32 students, we are also able to generalize the findings to patterns of preferences and effectiveness.CollaborationThis study concerns early experiments, starting with very little, to non-existent knowledge. As such, a sense of community is essential, yet the willingness of students to edit knowledge is not tested.

As such, a sense of community is essential, yet the willingness of students to edit knowledge is not tested. The incentives of sharing knowledge on manufacturing among peers are not yet explored. Similar phenomena, such as webpages of FabLabs should be studied as well – these are typically maintained by members to describe case studies and skills. Aspects of collaborative learning and Communities in Practice [13] theories should be included in further studies. Current examples such as the wiki of RepRap.org shows that facilitating discussions among members of a community is essential. Furthermore, educational settings are different from personal use of wikis, i.e.

Furthermore, educational settings are different from personal use of wikis, i.e. the motivation to add, reorganize and alter information is linked to grading the student’s assessment and not to social appreciation [14].Wiki structure and moderationOur objective is to establish an intertwined web of knowledge, with a several ways to navigate from hypothesis to solution, downloadable script/model and case studies. From the content of the wiki it is apparent that the participants are used to write reports of work done in chronological order. The documentation of just the results/data seems unfamiliar for the students. Overall these factors have made the content of the wiki less organized, but not necessarily incorrect. It has also made the wiki less expandable and not optimal for adding new knowledge. This is a typical challenge of wiki building and maintenance, some claim a hierarchical top- down structure such as a taxonomy is essential [15].

This is a typical challenge of wiki building and maintenance, some claim a hierarchical top- down structure such as a taxonomy is essential [15]. In this experiment, no active moderation of the pages was executed. As the experiment was done with a relatively small group, and took place within a short period of time, the entire wiki can still be overseen. It is clear that as the wiki would grow, structure and expandability will become increasingly important. An example of a database with active moderation is Thingiverse, this website offers ability to share 3D models in various formats, while designs are tracked but we need also for new ways of representing this information/kind of information. It includes a form of active moderation, e.g. to feature models on a front page.Conclusion and RecommendationAt present, designers are confronted with a shortage of knowledge to make effective use of the possibilities of Additive Manufacturing.

This article describes our efforts to understand how a DfAM method should be developed. In developing such a DfAM method we are facing several challenges: both AM technologies and supporting software are developing rapidly, posing the risk of any DfAM method being incomplete, or becoming obsolete within months. Furthermore, the possibilities of the currently existing technologies are still not fully explored and documented.We employed a wiki environment in an undergraduate prototyping course as a pilot study for open- ended knowledge management for DfAM. In four consecutive teams, 32 students have experimentally explored the manipulation of visual properties of AM by modeling and building novel AM structures. Each team utilized the knowledge on the wiki and in turn documented their findings for future groups. As the data on the wiki expanded, it encompassed knowledge on different levels, such as procedures and artifact descriptions.

As the data on the wiki expanded, it encompassed knowledge on different levels, such as procedures and artifact descriptions. The physical results showed a growth of sophistication, from simple tests to several products with remarkable visual properties.899We argue that building a DfAM method solely on currently available knowledge of the possibilities of AM is insufficient since AM clearly has many possibilities that are still unexplored. Based on this pilot study, we suggest several recommendations for the DfAM approach: it must offer a platform to allow mapping new knowledge on the possibilities of AM, and allow expansion of this knowledge triggered by the rapidly evolving possibilities of AM. Design acts, which were included in the pilot study, are proving to be a valuable source of such knowledge. Future DfAM methods should also address the organizations of collaboration and moderation, as well as conscious structuring of knowledge elements.References[1]S. Maidin, R. I.

Future DfAM methods should also address the organizations of collaboration and moderation, as well as conscious structuring of knowledge elements.References[1]S. Maidin, R. I. Campbell, I. Drstvensek, and P. Sever, “Design for Rapid Manufacturing – Capturing Designers Knowledge,” in Proceedings of the 4th International Conference on Advanced Research in Virtual and Rapid Prototyping, 2009, pp. 311–317.[2]I. Horváth, “Conceptual Design: Inside and Outside,” in Proceedings of the 2nd International Seminar and Workshop on Engineering Design in Integrated Product, 2000, pp. 63–72.[3] B. Singh and N. Sewell, “Knowledge based process planning and design for Additive Manufacturing ( KARMA ),” in Proceedings of the 5th International Conference on Advanced Research in Virtual and Rapid Prototyping, 2011, pp. 619–624.[4]J. Munguía, J. D. Ciurana, and C. Riba, “Pursuing successful rapid manufacturing: a users’ best-practices approach,” Rapid Prototyping Journal, vol. 14, no.

ve-Axis Freeform Fabrication of Thermoplastic Parts via SWIFT Tushar R. Mahale, James B. Taylor, Denis R. Cormier Department of Industrial Engineering, North Carolina State University, Raleigh, NC 27695AbstractCurrent Rapid Prototyping (RP) & Layered Manufacturing (LM) techniques are based on controlled, precise additive manufacturing methods. These techniques suffer from problems associated with stair stepping, loss of accuracy in the z-direction due to limitations on the slice thickness and their build speeds which prevent them from being “rapid”. Proposed here, is a method, which utilizes a freeform fabrication technique, coupled with the advantages and precision of five axis machining for contoured edge generation. A new technique for the generation of solid thermoplastic parts (ABS, Polystyrene, etc) will be introduced.

A new technique for the generation of solid thermoplastic parts (ABS, Polystyrene, etc) will be introduced. The proposed process will incorporate adaptive slicing for the generation of five-axis CNC code, and a sub-layer machining technique for use of uniform layer thickness will be outlined. A novel 5-axis configuration will be proposed. The proposed process will decrease, drastically, the time required for part production when compared to current commercial technologies.IntroductionRapid Prototyping (RP) has become synonymous with precision, layered manufacturing (LM). Though a fast method for developing prototypes, most of the commercially available processes have relatively high costs associated with them.

Though a fast method for developing prototypes, most of the commercially available processes have relatively high costs associated with them. These costs account for the fixed cost of equipment, and equally high variable cost of spares and consumables.LM processes are currently evolving from general rapid prototyping techniques to processes that provide industry specific manufacturing solutions (Karunakarn, 1999). Current markets have been exhibiting trends toward mass customization and LM will certainly have a place in the market.

Current markets have been exhibiting trends toward mass customization and LM will certainly have a place in the market. In spite of all the advances, almost all RP & LM processes still suffer from certain inherent deficiencies such as, stair stepping, low accuracy in comparison to machined parts, poor surface finish before post-processing, limitations on build materials, directional structural properties, and a non-homogeneous nature of the manufactured parts (Mahale, 1999, Tata, 1998).SWIFTA new process, Solvent Welding Freeform Fabrication Technique (SWIFT) involves the solvent welding of thermoplastic layers and the machining of layer contours (Taylor, 2000a). The process is being developed to meet the demands for generating low cost, high precision large to very large parts having fairly complex details.

The process is being developed to meet the demands for generating low cost, high precision large to very large parts having fairly complex details. An attempt has been made to counter some of the limitations of layered manufacturing processes by addressing speed, accuracy, strength and cost issues.In SWIFT, solvent weldable thermoplastic sheets of uniform layer thickness (sheets with variable layer thickness can also be used) are cut into a machine loadable size (currently a289maximum size of 7 inch x 10 inch can be accepted by the prototype machine). A suitable mechanism is used to automatically load the sheets onto the machine build platform. As the sheet is being passed through the loading mechanism, a uniform layer of solvent is applied to the bottom of the sheet. The entire bottom surface of the sheet is coated with the solvent when the sheet is brought over the build platform and then pressed over the previous layer by means of a platen.

The entire bottom surface of the sheet is coated with the solvent when the sheet is brought over the build platform and then pressed over the previous layer by means of a platen. The platen is held over the sheet in the same position for about 5-10 seconds to allow chemical bonds to form at the interface of the sheets, after which it is retracted. This assures a solid welding between the two sheets and the evaporation or absorption of the excessive solvent. On areas where bonding is not desired, a solvent mask can be applied. This can be the case especially on horizontal part surfaces. Once the platen is retracted, a face mill sizes the sheet to the required layer thickness. The face mill is retracted, and an end mill is loaded onto the spindle. At this stage, two more axes of control of the SWIFT machine are utilized (5-axis machining) to achieve the desired sloping surfaces on the edges of the contours of slices. The end mill machines the edge contours for the current slice.

The end mill machines the edge contours for the current slice. The whole process is repeated for the remaining layers. The SWIFT part is obtained by removing the build block from the build platform followed by the removal of the support structures. An alternative method for support stock removal would be required in the case of parts exhibiting a “bird cage” structure.SWIFT Machine ConfigurationConventional 5-axis machining centers consist of fully articulated heads i.e. two rotary axis are mounted on the spindle head and the three linear axes are distributed to the machine table and spindle head. LM involves the machining of one single, relatively thin layer at a time, and this layer is always positioned at the top of the part to be machined. Thus the cutting position of the bottom face of an end mill for cutting the edge contours would remain constant in the z-direction.

Thus the cutting position of the bottom face of an end mill for cutting the edge contours would remain constant in the z-direction. This suggests the possibility of utilizing only four simultaneously controlled axes during the process of edge contoured milling instead of the simultaneous 5-axis control as used in conventional 5-axis machining techniques (Ng, 1998, Wang, 1999). Such a machine would be a 4- _-axis machine rather than 5-axis.The proposed machine configuration calls for the build platform to be mounted on a rotary table so as to attain rotary motion along the Z-axis (C-axis). The rotary table would also have a linear motion along the Z-axis for between-layer positioning. The linear motion along the X & Y axes can be achieved through the standard motion of the spindle head. A unique configuration for the 5th axis (A-axis) has been proposed to specifically cater to the needs of LM processes.

A unique configuration for the 5th axis (A-axis) has been proposed to specifically cater to the needs of LM processes. The rotary motion of the A-axis is attained by pivoting the spindle about the Y axis, with the center of rotation coinciding with the bottom surface of the layer that is being machined. This greatly reduces the computational overhead associated with conventional 5-axis machining where the A-axis is pivoted above the cutting tool requiring the calculation of the simultaneous X, Y, and Z movement to be imparted to the table by way of compensation. It should be noted, the assembly for generating linear traverse along the y-axis is mounted on the A-axis, thus any movement to the A-axis results in a similar motion.290As seen in the figure below, due to the finite diameter of the tool, the 4-_-axis configuration could result in the previous layer being gouged.

The nature of pivoting for the A and C-axis would also limit the tool from completely cutting the down facing surfaces. A relation between the permitted deviation in the part dimensions, layer thickness, cutter diameter and the angle of the A-axis can help generate acceptable parts. Considering these limitations, the proposed configuration would be effective primarily for the manufacture of parts using large layer thickness (preferably using a small cutter diameter).> 90< 90SWIFT ConfigurationSWIFT ConfigurationGouged Region Uncut Region Figure 1 – Gouged and uncut regions in contoured edge generationAn alternative 5-axis configuration might partially help to solve the problem presented above. The 4-_-axis configuration presented above involves a setup where the A-axis moves tangential to the bottom surface of the machined layer.

The 4-_-axis configuration presented above involves a setup where the A-axis moves tangential to the bottom surface of the machined layer. In an alternative 5-axis configuration, the axis of rotation of the A-axis is moved to the intersection of the bottom surface of the tool and the tool’s axis of rotation. The proposed 5-axis configuration eliminates uncut regions, but increases the gouged region during the machining of down facing facets. The possibility of cutting into the previous layer while machining down facing part edges exists when the previous layer extends beyond the bottom of the top layer. Though a better configuration, it would be difficult to machine downward facing facets which are at a large angle with the z-axis. The figure below shows the proposed 5-axis configuration being utilized to machine upward facing facets. The face of the tool is used to machine facets where the angle made by the facet with the z-axis is less than 45 degrees.

The face of the tool is used to machine facets where the angle made by the facet with the z-axis is less than 45 degrees. In case where the angle between the facet and the z-axis is greater than 45 degrees, the bottom surface of the tool is used for machining the contoured edge. As is seen in the figure, this technique eliminates the gouged regions. It can be noted that in spite of utilizing simultaneous control of 5 axes, the amount of compensatory motion along the z-axis while machining a single layer is minimized.291Machining down facing facets 5-Axis ConfigurationA-axis degreestilt between 0-45A-axis tilt > 45 degreesFigure 2 – Alternate Edge contour generationA-axisC-AxisZ-Axis Y-AxisX-Axis C-Axis A-AxisTop View Front ViewFigure 3 – New 5-axis machine configurationAdaptive Slicing of STLThe name Layered Manufacturing suggests manufacturing by means of stacking laminated layers (Dolenc, 1995).

To generate the necessary data to conduct the LM operation, the CAD model of a part is intersected with parallel planes. The details generated through the intersection of a plane with the part contribute to the topographical details for that layer (Kulkarni, 1997). The operating cost of LM processes is a function of processing time, and the processing time is largely dependent on the number of layers required to build a part. One of the methods of decreasing the operation cost is to decrease the number of layers and hence the time required for manufacturing a part. The decrease in the number of layers is possible only by increasing the slice thickness. An increase in slice thickness (uniform slice thickness) over the entire part may result in the loss of accuracy along the Z-direction and substantial increase in stair stepping. The accuracy of a part can be maintained by adaptively selecting the layer thickness for individual layers as the part is being built.

The accuracy of a part can be maintained by adaptively selecting the layer thickness for individual layers as the part is being built. Though an effective method, zeroth order adaptive slicing can reduce but not eliminate instances of stair stepping. First order adaptive slicing (i.e. adaptively slicing the part for the generation of sloping surfaces from edge boundaries) is one of the best alternatives for approximating the surface of the STL model; thus minimizing stair stepping.Sub-Layer Machining292The SWIFT process involves the generation of layers through contour machining, thus the processing time per layer is related to the cutting speed and the length of the perimeter of the layer, as well as the overhead to apply the layer and size it. Face milling and sizing is conducted for each layer to maintain the accuracy of the part in the Z-direction as well as to assure a clean surface at the interface between two layers.

Face milling and sizing is conducted for each layer to maintain the accuracy of the part in the Z-direction as well as to assure a clean surface at the interface between two layers. Thin layers generated after adaptively slicing a part involve large cuts during the sizing, leading to the waste of raw material and an increase in processing time. An additional enhancement to this process would be to recognize the possibility of generating multiple thin layers from a base stock of higher thickness. This could play a big role in reducing the processing time, by greatly reducing the overhead time required for a group of sub-layers. In essence, the process of sub-layer machining treats a single thick layer of thermoplastic sheet to be composed of multiple layers of the sliced part. The result is that the time required for face milling (g ) for the ‘n’ sub-layers within the single thermoplastic sheet is reduced to g from the original n*g.

The result is that the time required for face milling (g ) for the ‘n’ sub-layers within the single thermoplastic sheet is reduced to g from the original n*g . Sub-layer machining is possible for layers that exhibit all up facing facets, i.e. layers consisting of facets whose k component of the normal is greater than zero (the capability of slicing software is currently limited to slicing along the Z-direction). This is simply because of the fact that it would be impossible to cut a bottom most sub-layer within a thermoplastic sheet without cutting into adjoining sub-layers.Up FacingDown FacingFigure 4 – Up and down facing sub-layer 3-axis edge contoursVolumetric Adaptive SlicingIn a 5-axis process, the use of 1st order adaptive slicing leaves us with the possibility of the occurrence of only the layering error (Hope, 1997a, Hope, 1997b, DeJager, 1996).

Layering error can be measured in terms of volume, Y , as it represents the region of the layer which has not been accurately generated i.e. it represents an uncut or gouged region.Cusp HeightLayering ErrorCusp HeightLayering Error1 1Figure 5 – Layering error in 5-axis edge generation293Earlier attempts at measuring layering error (Taylor, 2000b, Suh, 1994), involved approximating the volume Y , or by trying to determine the maximum difference in the layer plane, e . The value of e is considered to be a measure of layering error whereas the cusp height was considered to be a representation of the surface roughness. These measures are helpful when using the 0th order adaptive slicing. Layering error will not be exhibited when using 1st order adaptive slicing to slice a highly slanted planar surface, in spite of a large value of e.

Layering error will not be exhibited when using 1st order adaptive slicing to slice a highly slanted planar surface, in spite of a large value of e . The 1st order adaptive slicing technique might result in a non-zero cusp height; but this may not necessarily be a measure of surface roughness, as contoured edge generation tends to smooth the surface and eliminate the stair stepping.by determining the cusp height, xAs the prime objective of adaptive slicing is to reduce the layering error, thus, one of the criteria for adaptive slicing should be to actually measure the layering error (volume). The process begins by creating the first two slices of the part, beginning bottom up.

A user defined look up table helps the system to decide upon the number of layers the given layer is to be subdivided into. A sub-layer machining operation coupled with the volumetric slicing technique could be of great help in reducing the processing time, while simultaneously maintaining part accuracy.ConclusionThis paper has presented some of the results of research and development activities regarding rapid prototyping in general and the SWIFT process in particular. Software and295hardware enhancements to the process have recently added 5-axis sub-layer slicing. Software for 1st order adaptive slicing and machining capabilities and software for 5-axis contoured edge slicing has been developed and tested in a virtual environment. This enables the system to produce adaptively sliced, contoured edge parts using a uniform sheet thickness for the build material.

This enables the system to produce adaptively sliced, contoured edge parts using a uniform sheet thickness for the build material. A new configuration for a desktop 5-axis milling machine specifically designed for the SWIFT application has been proposed here and is currently under development. Future work will include building and testing the 5-axis machine and software enhancements to account for more difficult-to-build parts that include features such as “floating islands”, further reduction of processing time and other special features of historical importance to rapid prototyping development. On a 3-axis test bed, the process has proven its capability of producing, high strength thermoplastic parts. The SWIFT process has demonstrated advantages in low fixed and variable costs, part strength, part accuracy, process speed, and adaptability to advanced methods.

The SWIFT process has demonstrated advantages in low fixed and variable costs, part strength, part accuracy, process speed, and adaptability to advanced methods. At the present time the process is limited to thermoplastic parts and a rotational cutter of finite diameter serves as a limit to fine feature replication (though cutters of 0.031” have been successfully used).Figure 7 – SWIFT part with sublayer 3-axis machining detail, sublayer thickness of 0.012”.ReferencesDolenc, A., Makela, I., “Slicing Procedures for Layered Manufacturing Techniques”,Computer Aided Design, Vol. 26 No.2, 1995.De Jager, P. J., “Using Slanted and ruled layers for rapid prototyping”, in Dickens P. (Ed.), Proceedings of the 5th European Conference on Rapid Prototyping and Manufacturing, Helsinki, 1996.Hope, R.L., Jacobs, P.A. et al, “Rapid Prototyping with Sloping Surfaces”, RapidPrototyping Journal, Vol.3 No.1, 1997Hope, R.L., Roth, R.N.

AME RETARDANT INTUMESCENT POLYAMIDE 11 NANOCOMPOSITES – FURTHER STUDYS. C. Lao1,2, J. H. Koo1,2*, T. J. Moon1,2, W. Yong1, C. Lam1, J. Zhou1, B. Hadisujoto1, G. Wissler3, L. Pilato3, Z. P. Luo41The University of Texas at Austin, Department of Mechanical Engineering-C2200, Austin, TX 78712-02922The University of Texas at Austin, Texas Materials Institute 3KAI, LLC, Austin, TX 78739 4Texas A&M University *Corresponding author: jkoo@mail.utexas.eduAbstractThe objective of this research is to develop improved polyamide 11 and 12 polymers with enhanced flame retardancy, thermal, and mechanical properties for selective laser sintering (SLS) rapid manufacturing (RM). In the present study, a nanophase was introduced into the polyamide 11 and combine with a conventional intumescent flame retardant (FR) additive via twin screw extrusion.

In the present study, a nanophase was introduced into the polyamide 11 and combine with a conventional intumescent flame retardant (FR) additive via twin screw extrusion. Arkema Rilsan® polyamide 11 molding polymer pellets were used with two types of nanoparticles such as: chemically modified montmorillonite (MMT) organoclays and carbon nanofibers (CNFs). Two types of Clariant’s Exolit® OP 1311 and 1312 intumescent FR additives were used to generate a family of FR intumescent polyamide 11 nanocomposites with anticipated synergism.Transmission electron microscopy (TEM) and scanning electron microscopy (SEM) were used to determine the degree of dispersion. Injection molded test specimens were fabricated for physical, thermal, mechanical properties, and flammability measurements. Thermal stability of these intumescent polyamide 11 nanocomposites was examined by TGA.

Thermal stability of these intumescent polyamide 11 nanocomposites was examined by TGA. Mechanical properties such as ultimate tensile strength, rupture tensile strength, and elongation at rupture were measured. Flammability properties were also obtained using the UL 94 test method. TEM and SEM micrographs, physical, mechanical, and flammability properties are presented in the paper. FR intumescent polyamide 11 nanocomposites properties are compared with polyamide 11 baseline polymers. Based on flammability and mechanical material performance, selective polymers including polyamide 11 nanocomposites and control polyamide 11 will be cryogenically ground into fine powders for SLS RM processing. SLS specimens will be fabricated for flammability, mechanical and thermal properties characterization.1. IntroductionPolymeric resins are versatile materials and are used in a multitude of diverse areas from film, fabric to molded components for mechanical parts.

IntroductionPolymeric resins are versatile materials and are used in a multitude of diverse areas from film, fabric to molded components for mechanical parts. Inorganic fillers such as glass fibers, mineral fillers, and metallic fillers can be incorporated into these materials to further enhance their physical and mechanical properties such as toughness, stiffness, tensile strength, chemical resistivity, etc. Although physical and mechanical properties are enhanced by fillers, these resulting materials are lacking in flame retardancy and are of limited use in consumer productsthat require flame retardancy. Reduced flammability of these materials can be achieved by incorporating conventional fire retardant additives into these polymers.Flame retardant (FR) additives such as inorganic metal oxides/hydroxides or halogens with or without phosphorous and nitrogen containing materials are required in conventional methods to modify flammable thermoplastic materials as FR products [1].

Large amounts of FR additives (>30%) are necessary when those methods are used to make FR thermoplastics. In many cases a reduction of mechanical properties, such as toughness, melt flow, and/or release of smoke and toxic emissions, occurs when the modified thermoplastic is burning.The incorporation of nanoparticles has been shown to be an effective method for developing FR thermoplastic polymer by twin-screw extrusion (a melt blending process). It has been demonstrated that small amounts of nanoparticles (<7%) are required to make nanocomposites to exhibit similar enhanced flame retardant properties when compared with the FR thermoplastic generated by conventional methods [2-13]. However, unlike the conventional FR thermoplastics, resulting nanocomposites exhibit enhanced mechanical properties such as high the strength/modulus, moisture resistance, higher heat deflection temperature, etc.

Therefore, nanotechnology can be used to develop novel FR thermoplastic structural components with high performance characteristics.Our previous studies [14-15] indicated that low amounts (5%) of nanoparticles enhance the mechanical properties (such as tensile strength, flexural strength, and Young’s modulus) but are somewhat lacking in flame retardancy to those FR thermoplastics with conventional FR additives. Large amounts of nanoparticles (>7%) would yield relatively good enhancement of flame retardancy, but the mechanical properties of the resulting polymer nanocomposites would be compromised. Therefore, the preparation of a truly fire retardant PA11, requires the combination of the conventional FR additives and nanoparticles together into the polymer resin.

Therefore, the preparation of a truly fire retardant PA11, requires the combination of the conventional FR additives and nanoparticles together into the polymer resin. It is expected that if synergy between the FR additives and nanoparticles occurs within the PA11 resin, then the flammability and thermal properties of the resulting PA11 FR nanocomposites would be improved as compared with PA11 nanocomposite or PA11 solely with FR additives. Our related study that involved PA11 FR nanocomposites [16] demonstrated that synergism existed between the Clariant’s OP1230 FR additives and Southern Clay Products’ Cloisite 30B clay/Applied Science’s Pyrograf® PR-19-PS CNF and superior thermal and flammability properties were observed as anticipated. Furthermore, the mechanical properties (except low tensile elongation at rupture) of these PA11 FR nanocomposites were improved when compared to those of the PA11 with only the FR additives.

Furthermore, the mechanical properties (except low tensile elongation at rupture) of these PA11 FR nanocomposites were improved when compared to those of the PA11 with only the FR additives. However, PA11 FR nanocomposites were still deficient to those of the neat PA11 or the PA11 nanocomposites [17].Thus the major technical objective of this research program is to further enhance the mechanical properties of these PA11 FR nanocomposites by using additional, selective FR additives while maintaining the same amount of nanoclay and CNF. It is anticipated if synergism occurs between the nanoparticles and the conventional FR additives in the PA11 resin, a reduced amount of FR additives will be needed to prepare the new formulations to pass the flammability test with enhanced mechanical properties. Various weight loadings of nanoclay (5 to 7.5%), carbon nanofiber (5 to 7.5%), and the newly selected intumescent FR (15 to 25%) were melt- blended with polyamide 11 via twin-screw extrusion.

Various weight loadings of nanoclay (5 to 7.5%), carbon nanofiber (5 to 7.5%), and the newly selected intumescent FR (15 to 25%) were melt- blended with polyamide 11 via twin-screw extrusion. The resulting pellets were injection molded into different test specimens.Transmission electron microscopy (TEM) was used to examine the degree of dispersion of nanoparticles. Initial flammability properties were studied by conducting the UL 94 test. Thermal properties of the samples were analyzed using thermogravimetric analysis (TGA). Mechanical properties were measured in accordance with the ASTM D638. Results were compared with those of FR PA11 nanocomposites with OP1230 that were reported previously [16-17].2.

Mechanical properties were measured in accordance with the ASTM D638. Results were compared with those of FR PA11 nanocomposites with OP1230 that were reported previously [16-17].2. Experimental2.1 MaterialsPolymer Resin Polyamides (PA) are versatile engineering thermoplastics noted for outstanding properties such as high tensile strength, good resistance to flow under pressure (creep), excellent abrasion, chemical and heat resistance, and a low coefficient of friction. Polyamides such as PA6 and PA66 are high melting, moderately crystalline polymers (Tm 220- 265°C) while PA11 and PA12 are intermediate melting materials (<200°C) and are less susceptible to moisture as compared to PA6 and PA66. Paul et al. [18] have examined the structure and properties of nanocomposites based on PA6, PA11, and PA12.

Paul et al. [18] have examined the structure and properties of nanocomposites based on PA6, PA11, and PA12. Their studies were focused solely on nanoclays whereas our studies are directed to examining/comparing nanoclay and other nanoparticles to determine enhanced polymer characteristics such as flame retardancy and improved thermal/mechanical properties for the resulting PA11N.Arkema’s RILSAN® polyamide 11 (PA11) was selected for this study since it is an attractive polyamide used in a variety of applications. RILSAN® PA11 thermoplastic [19] is a polymer developed by Atofina Chemicals, Inc. (now known as Arkema, Inc.) in 1942. Derived from a series of complex chemical operations, RILSAN® PA11 is one of the few polymers that is produced from ‘green’ raw materials – castor beans. RILSAN® PA11 resin has earned a preferred material status in the most demanding applications due largely to their unique combination of thermal, physical, chemical, and mechanical properties.

RILSAN® PA11 resin has earned a preferred material status in the most demanding applications due largely to their unique combination of thermal, physical, chemical, and mechanical properties. This results in an outstanding "cost performance ratio." Processing ease is another major benefit of RILSAN® polyamide 11 resins. Supplied in powder or pellet form RILSAN® PA11 resin can be processed by injection molding, extrusion, blown film extrusion, extrusion blow molding or rotomolding. The ease of processing of Rilsan® has led designers to select them for industries as diverse as aerospace, offshore drilling, electrical cables, automotive, and pneumatic and hydraulic hose.Nanoparticles Two types of nanoparticles, nanoclay and carbon nanofiber, were used. Southern Clay Products’ montmorillonite (MMT) nanoclay, Cloisite® 30B (a natural MT modified with an organic modifier, MT2EtOT: methyl-tallow-bis-2-hydroxyethylquaternary ammonium at 90 meq/100g) was selected [20].

Achieving exfoliation of organomontmorillonite in various continuous phases is a function of the surface treatment of the MMT clays and the mixing efficiency of the dispersing equipment. Surface treatment of MMT is conveniently accomplished with the exchange of inorganic counterions, e.g., sodium etc., with quaternary ammonium ions. The second type of nanoparticles used was Pyrograf®-III PR-19-PS carbon nanofiber (CNF) from Applied Science, Inc. CNF is a form of vapor-grown carbon fiber, which is a discontinuous graphitic filament produced in the gas phase from the pyrolysis of hydrocarbons [21-24].

CNF is a form of vapor-grown carbon fiber, which is a discontinuous graphitic filament produced in the gas phase from the pyrolysis of hydrocarbons [21-24]. In regard to properties of physical size, performance improvement, and product cost, CNF complete a continuum bounded by carbon black, fullerenes, and single-wall to multi-wall carbon nanotubes on one end and continuous carbon fiber on the other end [24].Fire Retardant Additive Clariant’s Exolit® OP1311 and OP1312, which are formulated for polyamide materials, are non-halogentaed intumescent flame retardant additives based on organic phosphinate type materials. They are non-toxic, hydrophobic, and exhibit good thermal stability.

They are non-toxic, hydrophobic, and exhibit good thermal stability. Exolit materials were mixed with polyamide 66, and the resulting products showed enhanced FR properties [25-27].2.2 MeasurementsMorphological Microstructures Analysis The uniformity of the PA11 FR nanocomposites was investigated by TEM to examine the dispersion of nanoclay, carbon nanofibers, and intumescent FR additives within the PA11 polymer matrix. Uniform distribution of the nanoparticles/FR additives within the polymer matrix is essential to yield the best enhancement of material properties of the polymer matrix.Mechanical Testing In order to compare the mechanical properties of nanocomposites, stress-strain behavior was performed using an automated tensile testing system (Instron model 3345) in accordance with ASTM D638.

Ultimate tensile strength, rupture tensile strength, tensile modulus, and tensile elongation at rupture were obtained from the stress-strain data.Thermal Stability Testing Thermal stability of the PA11 baseline and PA11N were examined by thermogravimetric analysis (TGA) using the Perkin Elmer TGA 7. Weight changes in sample materials are measured as a function of temperature or time in TGA. The sample is heated by a furnace with nitrogen while the loss or gain of sample weight is monitored by a sensitive balance. Weight, temperature, and furnace calibrations were carried out within the range of the TGA (100-900°C) at scan rates of 10°C/min and 20°C/min.Flammability Testing UL 94, a standard test for flammability of plastic materials in industry that serves as a preliminary indication of plastics acceptability for use as a component within a device or appliance with respect to its flammability behavior.

UL 94 is not intended to reflect the hazards of a material under actual fire conditions but is considered as a preliminary step toward obtaining plastic recognition and subsequent listing in the “Plastics Recognized Component Directory” (formerly known as “Yellow Cards”). The materials are tested in a vertical setting to determine the UL94 V-0 rating. UL 94 Vertical burning test involves a 1/2” x 5” specimen which is held at one end in the vertical position. A burner flame is applied to the free end of the specimen for two 10 second intervals separated by the time it takes for flaming combustion to cease after the first application. Five specimens are tested for each formulation.3. Results and Discussion3.1 Processing and Characterization of the Polymer NanocompositesBlending Nanoclays/FR Additive/PA11 Polymer Chemically treated pillared clay Cloisite® 30B and intumescent fire retardant additive Exolit® OP1311 and 1312 were used.

The individual clay layers are separated by alkyl ammonium ion incorporation (d spacing increased) allowing for anticipated intercalation of solid organic resins when the clay is melt blended with the resins. Clays were blended with the PA11 resin to intercalate and eventually exfoliate the clay. Twin screw extrusion mixing enhances the exfoliation rate. The degree of exfoliation is examined by TEM.A 30 mm Werner Pfleider corotating twin screw extruder which is configured for a widevariety of materials was used. The extruder L/D can be varied from 21 to 48, with options of multiple feeds and vents. The energy profile of the screw is adjusted to optimally meet the needs of the target product. Table 1 shows the nanoparticles and FR additives selected at various wt% loading levels with the PA11 resin.

Table 1 shows the nanoparticles and FR additives selected at various wt% loading levels with the PA11 resin. For the clay a relatively long residence time is preferred for the screw profile allowing high shear characteristics for complete separation of the nanoclay platelets so that the polymer matrix is dispersed uniformly into the clay galleries. Approximately 10 lbs of each formulation were produced. Separate volumetric feeders were used for the base resin and the nanoparticles. The PA11 was dried in a desiccant drier before compounding. Injection molded specimens of each blend were prepared and examined by TEM.Figure 1 shows the TEM micrographs of polyamide 11 with 15% OP1311 and 5% Cloisite® 30B. It is evident that exfoliation of Cloisite® 30B in PA11 polymer was achieved.

It is evident that exfoliation of Cloisite® 30B in PA11 polymer was achieved. The micro- sized FR intumescent FR additives were also uniformly dispersed in the PA11 polymer.Table 1 Composition of various formulations of FR PA11N and UL 94 resultsBlending CNF/FR Additive/PA11 Polymer PR-19-PS CNF possesses a diameter of about 130 nm and a length of several microns. It can be classified as MWNT. It was blended with PA11 polymer along with FR additive in three different loading levels via twin screw extrusion (Table 1). For the CNF, the screw was configured to allow separation of the CNF without fracturing them. Figure 2 shows the TEM micrographs of polyamide 11 with 15% OP1311 and 5% CNF.

For the CNF, the screw was configured to allow separation of the CNF without fracturing them. Figure 2 shows the TEM micrographs of polyamide 11 with 15% OP1311 and 5% CNF. The CNF was also well dispersed in the matrix.Figure 1 TEM micrographs of polyamide 11 with 15% OP1311 (large dark micron-sized particles) and 5% Cloisite® 30B (hairline 1 nm thick platelets) polymer showing exfoliation of nanoclay in PA11 was achieved.Figure 2 TEM micrographs of polyamide 11 with 15% OP1311 (dark micron-sized particles) and 5% PR-19-PS CNF (elongated fibers).3.2 Mechanical Properties of the Polymer NanocompositesUltimate tensile strength, rupture tensile strength, tensile modulus, and tensile elongation at rupture of the PA11 FR nanocomposites were measured (Figures 3-6). In all cases beside tensile modulus, the PA11 FR nanocomposites with 5%CNF had better properties than those with5%30B possessing the same type and loading level of FR additives.

In all cases beside tensile modulus, the PA11 FR nanocomposites with 5%CNF had better properties than those with5%30B possessing the same type and loading level of FR additives. At a constant amount of 30B and combined with OP1312, that composition yielded the best ultimate tensile strength and rupture tensile strength, while for CNF, OP1230 gave the best results.Figure 3 Ultimate tensile strength of PA11 FR nanocomposites with 5% nanoparticles.Figure 4 Rupture tensile strength of PA11 FR nanocomposites with 5% nanoparticles.Figure 5 Tensile modulus of PA11 FR nanocomposites with 5% nanoparticles.Figure 6 Tensile elongation at rupture of PA11 FR nanocomposites with 5% nanoparticles.3.3 Thermal Stability of the Polymer NanocompositesTGA was performed on all formulations under nitrogen using the heating rate of 20C/min (Figures 7-9).

In all cases, PA11 FR nanocomposites had better thermal stability or higher decomposition temperatures than the neat PA11 or the related solely PA11 nanocomposites. Furthermore, PA11 FR nanocomposites with 5% 30B had better thermal stability than those with 5% CNF and same type and loading level of FR additives.

Furthermore, PA11 FR nanocomposites with 5% 30B had better thermal stability than those with 5% CNF and same type and loading level of FR additives. At the loading of 5% 30B and combined with OP1230, the resulting composition exhibited the best thermal stability, while for CNF combined with OP1230, it possessed the highest decomposition temperature at 10% mass loss but OP1312 provided the highest decomposition temperature at 50% mass loss.Figure 7 Thermogravimetric analysis of PA11 FR nanocomposites with 20% FR and 5% nanoparticles at 20C/min in nitrogen.Figure 8 Thermogravimetric analysis of PA11 FR nanocomposites with 20% FR and Cloisite 30B at 20C/min in nitrogen.Figure 9 Thermogravimetric analysis of PA11 FR nanocomposites with 20% FR and CNF at 20C/min in nitrogen.3.4 Flammability Properties of the Polymer NanocompositesThe materials were tested as “received with no additional conditioning/drying” for UL 94 testing.

The test was performed in our lab and the UL 94 testing requirements. Procedures were followed as stringently as possible. Our lab is not certified for UL 94, and the results serve as a screening tool. Five specimens were tested for each formulation (except formulation #9 in Table 1 since specimens were not available). The testing was performed in a fume hood with a preset airflow of 90-105 ft/min. Fume hood sash was pulled down as much as possible to prevent airflow from the outside environment. The erratic exhaust airflow from the fume hood and the surrounding made the reproducibility of the testing challenging. The burner is lit during the time of this experiment to keep the applied flame constant between each specimen. Table 1 shows the UL 94 results. It showed that even though the FR additives were used together with nanoparticles, loadings of 10% and 15% of FR additives were insufficient for V0 rating.

Table 1 shows the UL 94 results. It showed that even though the FR additives were used together with nanoparticles, loadings of 10% and 15% of FR additives were insufficient for V0 rating. Moreover, of all the formulations with 5% 30B and 20% FR additives (#4, 13, and 21), only 30B combined with OP1312 achieved the UL94 V0 rating. In regard to CNF, all formulations with at least 20% FR additives and at least 5% CNF achieved UL94 V0 rating. It seems that CNF exhibited a better synergistic effect with all three types of FR additives than 30B as it related to flammability retardancy.4. Summary and ConclusionsA total of 26 polymer blends of polyamide 11 polymer with MMT nanoclay/carbon nanofiber and intumescent fire retardant additive were compounded separately via twin screw extrusion. Injection molded specimens of PA11 baseline and FR-PA11N were fabricated for physical, mechanical, thermal, and flammability properties measurements.

Injection molded specimens of PA11 baseline and FR-PA11N were fabricated for physical, mechanical, thermal, and flammability properties measurements. Some results were also compared with PA11 clay nanocomposites. Transmission electron microscopy was used to examine the dispersion uniformity of the nanoparticles in the polyamide 11 resin. Thermalstability of PA11N was analyzed using thermogravimetric analysis. Flammability properties were measured by UL 94 test.The following conclusions were drawn from this study:1. The TEM analysis has shown that good dispersions of either nanoclays or carbon nanofibers in the polyamide 11 resin systems were achieved.2. In regard to mechanical properties, OP1312 exhibited enhanced flame retardance with 30B while OP1230 performed better with CNF.3. In terms of thermal stability, OP1230 was best when combined with 30B while OP1312 was best with CNF.4.

In terms of thermal stability, OP1230 was best when combined with 30B while OP1312 was best with CNF.4. As it related to flammability properties, 30B combined best with OP1312 while CNF performed well with all three FR additives.5. In this study, a minimum of 20% FR additives is needed to pass the UL 94 V0 rating.6. The threshold amount of FR additives needed to achieve UL94 V0 rating can be lower (between 15% and 20%) but needs further investigation.7. A third element, e.g. elastomer, may be required to further improve toughness.8. Selected formulations will be chosen for SLS fabrication.5. References1. G.N. Nelson and C.A. Wilke, Ed., Fire and Polymers III. Materials and Solutions for Hazard Prevention, American Chemical Society, Washington, D.C., 2001.2. E. Giannelis, Adv. Mater. 8, 29 (1996). 3. J.W. Gilman, T. Kashiwagi, SAMPE J. 33 (4), 40 (1997). 4. J.W. Gilman, Appl. Clay. Sci. 15, 31 (1999). 5. J. Zhu, A.B. Morgan, J. Lamelas, C.A. Wilke, Chem. Mater.

AMMABILITY AND THERMAL PROPERTIES OF POLYAMIDE 11-ALUMINA NANOCOMPOSITESS. C. Lao1, J. H. Koo1*, T. J. Moon1, B. Hadisujoto1, W. Yong1, L. Pilato2, and G. Wissler2 1The University of Texas at Austin, Department of Mechanical Engineering-C2200, Austin, TX 78712-0292 2KAI, Inc., Austin, TX 78739Corresponding author: jkoo@mail.utexas.eduAbstractNeat polyamides 11 and 12 lack high strength/high heat resistance and flame retardancy. The incorporation of selected nanoparticles is expected to enhance these properties to a level that is desired and required for performance driven applications. This enhancement may result in additional market opportunities for the polyamide 11 and 12 polymer manufacturers. The objective of this study is to develop polyamide 11 polymer nanocomposites with enhanced thermal, flammability, and mechanical performance for selective laser sintering (SLS) rapid manufacturing.

The objective of this study is to develop polyamide 11 polymer nanocomposites with enhanced thermal, flammability, and mechanical performance for selective laser sintering (SLS) rapid manufacturing. Three types of nano-alumnia (X-0 needle, X-25SR, and X-0SR) with different organic treatments were melt-compounded into polyamide 11 in three different weight loadings of the nanoparticles (2.5%, 5%, and 7.5%). Injection molded specimens were fabricated for thermal, flammability, and mechanical properties characterization. Although nano-alumina was uniformly dispersed in polyamide 11 and better thermal stability of the nanomodified materials was observed, the desired FR characteristics of the nanomodified polyamide 11 was not achieved. None of the materials passed the desired UL 94 V0 rating.1.

None of the materials passed the desired UL 94 V0 rating.1. IntroductionFlame retardant (FR) additives such as inorganic metal oxides/hydroxides or halogens with or without phosphorous and nitrogen containing materials are required in conventional methods to modify flammable thermoplastic materials as FR products [1]. Large amounts of FR additives (>30%) are necessary when using these methods to make FR thermoplastics. In many cases a reduction of mechanical properties, such as toughness, melt flow, etc. and/or release of smoke and toxic emissions, occurs when the modified thermoplastic is burning.The incorporation of nanoparticles has been shown to be an effective approach for developing FR thermoplastic polymer by twin-screw extrusion (a melt blending process). Small amounts of nanoparticles (<7%) are required to prepare nanocomposites that the modified exhibit enhanced thermoplastic processed by conventional methods [2-13].

Small amounts of nanoparticles (<7%) are required to prepare nanocomposites that the modified exhibit enhanced thermoplastic processed by conventional methods [2-13]. However, unlike the conventional FR the resulting nanocomposites exhibit enhanced mechanical properties such as high strength/modulus, moisture resistance, higher heat deflection temperature, etc. Therefore, nanotechnology can be used to develop novel FR thermoplastic structural components with improved, high performance characteristics.flammability properties when compared withthermoplastics,The major technical objective of this study is to investigate the feasibility of enhancing the fire retardancy and flammability properties of nylon (polyamide) thermoplastic material by combining it with selected nanoparticles. It involved the incorporation of nanoparticles into nylon (polyamide) 11 (PA11) to form a nylon (polyamide) 11 nanocomposite (PA11N).

It involved the incorporation of nanoparticles into nylon (polyamide) 11 (PA11) to form a nylon (polyamide) 11 nanocomposite (PA11N). These PA11N materials are expected to exhibit reduced flammability and improved thermal properties as compared to the baseline PA11 material. The effect of nanoparticles on the resin will also be examined. Three types of nano- alumina (n-alumina) with different organic surface treatments and shape, namely Dispal® X-0 needle, X-25SR, and X-0SR, were melt-blended, separately. These n-alumina were compounded with RILSAN® PA11 using twin-screw extrusion to form PA11N in three different weight loadings (2.5%, 5%, and 7.5%). A total of ten formulations (including the neat PA11) were produced. The resulting pellets were injection molded into different test specimens. Transmission electron microscopy (TEM) was used to examine the extent of dispersion of nanoparticles in the polymer. Flammability properties were studied by conducting UL 94 test.

Transmission electron microscopy (TEM) was used to examine the extent of dispersion of nanoparticles in the polymer. Flammability properties were studied by conducting UL 94 test. Thermal properties of the samples were analyzed using thermogravimetric analysis (TGA).2. Experimental2.1 MaterialsPolymer Resin Polyamides (PAs) are versatile thermoplastic engineering polymeric materials noted for outstanding properties such as high tensile strength, good resistance to flow under pressure (creep), excellent abrasion, chemical and heat resistance, and a low coefficient of friction. Polyamides such as PA6 and PA66 are high melting, moderately crystalline polymers (Tm 220-265°C) while PA11 and PA12 are intermediate melting materials (<200°C) and are less susceptible to moisture as compared to PA6 and PA66. Fornes and Paul [14] have examined the structure and properties of nanocomposites based on PA6, PA11, and PA12.

Fornes and Paul [14] have examined the structure and properties of nanocomposites based on PA6, PA11, and PA12. Their studies were focused solely on nanoclays whereas our studies are directed to examining/comparing nanoclay and other nanoparticles to determine enhanced polymer characteristics such as flame retardancy and improved thermal and mechanical properties for the resulting PA11N.Arkema’s RILSAN® polyamide 11 (PA11) was selected for this study since it is an attractive polyamide used in a variety of applications. RILSAN® PA11 thermoplastic [15] is a polymer developed by Atofina Chemicals, Inc. (now known as Arkema, Inc.) in 1942. Derived from a series of complex chemical operations, RILSAN® PA11 is one of the few polymers that are produced from ‘green’ raw materials – castor beans. RILSAN® PA11 resin has earned a preferred material status in the most demanding applications due largely to their unique combination of thermal, physical, chemical, and mechanical properties.

RILSAN® PA11 resin has earned a preferred material status in the most demanding applications due largely to their unique combination of thermal, physical, chemical, and mechanical properties. This results in an outstanding "cost performance ratio." Processing ease is another major benefit of RILSAN® polyamide 11 resins. Supplied in powder or pellet form RILSAN® PA11 resin can be processed by injection molding, extrusion, blown film extrusion, extrusion blow molding or rotomolding. The ease of processing of Rilsan® has led designers to select them for industries as diverse as aerospace, offshore drilling, electrical cables, automotive, and pneumatic and hydraulic hose.Nanoparticles. Three types of n-alumina with different organic surface treatment and shape were used. They were Sasol’s Dispal® X-0 needle, X-0 SR, and X-25 SR. Theyare synthetic alumina nanopartilces which are high purity, highly dispersible, colloidal Boehmite alumina powders manufactured by Sasol.

Theyare synthetic alumina nanopartilces which are high purity, highly dispersible, colloidal Boehmite alumina powders manufactured by Sasol. These alumina powders are available in micron size but they become nano-sized in the dispersed phase. Figure 1 shows the high resolution SEM micrographs of as-received n-alumina.Figure 1 SEM micrographs of as-received n-alumina; X-0 needle (left), X0-SR (middle), and X25-SR (right).2.2 Measurementsthe PA11 Morphological Microstructures Analysis The cross-sections of nanocomposites were investigated by TEM to examine the dispersion of n-alumina within the PA11 polymer matrix. Uniform distribution of the nanoparticles within the polymer matrix is essential to yield the best enhancement of material properties of the polymer matrix.Thermal Stability Testing Thermal stability of the PA11 baseline and PA11N were examined by thermogravimetric analysis (TGA) using the Perkin Elmer TGA 7.

Weight changes in sample materials are measured as a function of temperature or time in TGA. The sample is heated by a furnace with nitrogen while the loss or gain of sample weight is monitored by a sensitive balance. Weight, temperature, and furnace calibrations were carried out within the range of the TGA (100-900°C) at scan rates of 10°C/min and 20°C/min.Flammability Testing UL 94 is a standard test for flammability of plastic materials in industry that serves as a preliminary indication of plastics acceptability for use as a component of a device or appliance with respect to its flammability behavior. UL 94 is not intended to reflect the hazards of a material under actual fire conditions but is considered as a preliminary step toward obtaining plastic recognition and subsequent listing in the “Plastics Recognized Component Directory” (formerly known as “Yellow Cards”). The materials are tested in a vertical setting to determine the UL 94 V-0 rating.

The materials are tested in a vertical setting to determine the UL 94 V-0 rating. UL 94 Vertical burning test involves a 1/2” x 5” specimen which is held at one end in the vertical position. A burner flame is applied to the free end of the specimen for two 10 second intervals separated by the time it takes for flaming combustion to cease after the first application. Five specimens are tested for each formulation.3. Results and Discussion3.1 Processing and Characterization of the Polymer NanocompositesA 30 mm Werner Pfleider corotating twin screw extruder which is configured for a widevariety of materials was used. The extruder L/D can be varied from 21 to 48, with options of multiple feeds and vents. The energy profile of the screw is adjusted to optimally meet the needs of the target product. Approximately 10 lbs of each formulation were produced. Separate volumetric feeders were used for the base resin and the nanoparticles.

Approximately 10 lbs of each formulation were produced. Separate volumetric feeders were used for the base resin and the nanoparticles. The PA11 was dried in a desiccant drier before compounding. Injection molded specimens of each blend were prepared and examined by TEM. Figures 2 to 4 show TEM micrographs of PA11 with 5% X-0 needle, 5% X0-SR, and 5% X25-SR, respectively, which are indication that good dispersion of the nanoparticles in the polyamide matrix were achieved.Figure 2 TEM micrographs of polyamide 11 with 5% X-0 needle n-alumina.Figure 3 TEM micrographs of polyamide 11 with 5% X0-SR n-alumina.Figure 4 TEM micrographs of polyamide 11 with 5% X25-SR n-alumina.3.2 Thermal Stability of the Polymer NanocompositesTGA was performed on the nine polymer blends and control under nitrogen using a scan rate of 10C/min. Table 1 shows the decomposition temperatures of the PA11 nanocomposites as well as the neat nylon 11.

Table 1 shows the decomposition temperatures of the PA11 nanocomposites as well as the neat nylon 11. All PA11 n-alumina nanocomposites had substantial increases in decomposition temperatures (from 26 to 46oC) compared to the neat nylon 11. There was not a clear trend of thermal properties enhancement among different types of n-alumina or different weight loading. However, the lowest weight loading of each type of n-alumina appeared to provide the best improvement in decomposition temperature. Figures 5 to 7 showed the TGA plots of nylon 11 and the nanocomposites with 2.5%, 5%, and 7.5%, respectively, of each type of the n-alumina. It has been clearly shown that in all loading levels of n-alumina, the PA11/n-alumina polymer nanocomposites have better thermal stability than the neat PA11 polymer.Table 1.

Decomposition Temperatures of PA11/n-Alumina NanocompositesSamplesDecomposition Temperature (C)ΔT (C)PA11 PA11/2.5% X-0 PA11/5.0% X-0 PA11/7.5% X-0 PA11/2.5% X0-SR PA11/5.0% X0-SR PA11/7.5% X0-SR PA11/2.5% X25-SR PA11/5.0% X25-SR PA11/7.5% X25-SR421 467 459 464 457 453 458 455 447 448--- 46 38 43 36 32 37 34 26 27100PA11 neat902.5% X-0 Needle802.5% X-0 SR702.5% X-25 SR60%t h g i e W50403020100 300350400450500550Temperature (C)Figure 5 TGA of polyamide 11 with 2.5% n-alumina at 10C/min in nitrogen.100PA11 neat905% X-0 Needle805% X-0 SR705% X-25 SR60%t h g i e w50403020100 300350400450500550Temperature (C)Figure 6 TGA of polyamide 11 with 5% n-alumina at 10C/min in nitrogen.600600100PA11 neat907.5% X-0 Needle807.5% X-0 SR707.5% X-25 SR60%t h g e Wi50403020100 300350400450500550Temperature (C)Figure 7 TGA of polyamide 11 with 7.5% n-alumina at 10C/min in nitrogen.3.3 Flammability Properties of the Polymer NanocompositesUL 94 The materials were tested as “received with no additional conditioning/drying” before UL 94 testing.

The test was performed in our lab and the UL 94 testing requirements and procedures were followed as stringently as possible. Our lab is not certified for UL 94 and the results serve as a screening tool. Five specimens were tested for each formulation. The testing was performed in a fume hood with a preset airflow of 90-105 ft/min. Fume hood sash was pulled down as much as possible to prevent airflow from the outside environment. The erratic exhaust airflow from the fume hood and the surrounding made the reproducibility of the testing challenging. The burner is lit during the time of this experiment to keep the applied flame constant between each specimen.As shown in Table 2, the addition of n-alumina to the PA11 material increased the FR properties of the nylon 11 by reducing dripping and flaming combustion time.

Though all formulations of the PA11N failed the V-0 and V-1 ratings, the X-0 needle formulations performed the best in terms of average burn time and resistance to dripping. The X-0 SR performed the worst, with several samples completely burned, and the X-25 SR was the middle performer.600Table 2. UL 94 Results of PA11/n-Alumina NanocompositesSamples1st Burn Drip2nd Burn DripDrip Rank (1 = best)UL 94 RatingPA11 PA11/2.5% X-0 PA11/5.0% X-0 PA11/7.5% X-0 PA11/2.5% X-0 SR PA11/5.0% X-0 SR PA11/7.5% X-0 SR PA11/2.5% X-25 SR PA11/5.0% X-25 SR PA11/7.5% X-25 SRYes Yes Yes Yes Yes Yes Yes Yes Yes YesYes Yes Yes Yes Yes Yes Yes Yes Yes Yes3 1 1 1 3 3 3 2 2 2Failed V-2 V-2 V-2 Failed Failed Failed V-2 Failed FailedNote: To obtain a UL 94 V-0 rating, one of the criteria conditions is that the afterflame time for each individual specimen t1 (afterflame time after first flame application) or t2 (afterflame time after second flame application) has to be ≤10s.

For V-1 rating the afterflame time for each individual specimen t1 or t2 has to be ≤30s.X-0 needle Overall these samples exhibited the highest resistance to dripping compared to the other PA11N formulations. On average, they also had the lowest total flaming combustion time. The 5% weight loading achieved the lowest total flaming combustion time. All three weight loadings achieved a V-2 rating.X-25 SR These samples exhibited less resistance to dripping than the X-0 needle, but performed comparably in terms of total flaming combustion time. Increasing the weight loading resulted in higher total flaming combustion times. This is because the drip resistance of the PA11N was increased, but allowed for flame propagation up the material.

Increasing the weight loading resulted in higher total flaming combustion times. This is because the drip resistance of the PA11N was increased, but allowed for flame propagation up the material. For the lower weight loadings, the flaming portion of the sample was allowed to drip and fall, not allowing enough time for the flame to propagate.X-0 SR The PA11 samples with X-0 SR exhibited high total combustion times and several of the samples burned completely. Increasing the weight loading increased the total combustion times, and several samples did not withstand the first 10 second flame application.4. Summary and ConclusionsA total of ten formulations of polyamide 11/n-alumina nanocomposites, including the neat PA11, were melt-compounded and injection molded to test specimens. TEM demonstrated that n-alumina surface treated particles are well dispersed in the polyamide 11 matrix. TGA showed that n-alumina improved the thermal stability of the nylon 11 substantially.

TEM demonstrated that n-alumina surface treated particles are well dispersed in the polyamide 11 matrix. TGA showed that n-alumina improved the thermal stability of the nylon 11 substantially. The UL 94 test showed that all formulations did not pass the ideal V-0 rating. All of them exhibited flammable dripping that ignited the cotton under the specimens during the test. The X-0 needle formulations performed the best, followed by X-25 SR, and the X-0 SR performed the worst in performance. The addition of selective flame retardant additives or anti-dripping agents will be incorporated into the systemtogether with the n-alumina for enhanced flammability properties. Additional tests, such as mechanical testing, cone calorimentry, heat deflection temperature, etc., will be carried out to fully characterize these polyamide 11/n-alumina nanocomposites so that they can be considered for use in consumer products. Additional research is ongoing.5. References1.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFlammability of 3-D Printed Polymers – Composition and Geometry FactorsStefan M. Kraft, Brian Y. Lattimer, Christopher B. Williams Department of Mechanical Engineering, Virginia TechAbstractThe focus of this paper is to evaluate the comparative flammability of additively manufactured (AM) and conventionally molded polymers. Flammability of objects is dependent on two main factors: material composition and object geometry. To evaluate effects of material composition, experiments on polymer samples made via conventional molding and via AM were performed using an ASTM E1354 cone calorimeter to measure and compare material ignitability and heat release rate.

thermoplastic polyetherimide) and PPSF/PPSU (polyphenylsulfone) heat release rates were about 10 times lower than ABS (acrylonitrile butadiene styrene). This was in part due to the large char layer formed by these materials during burning. Comparisons between conventional molded and AM materials revealed slight differences in heat release rate. Additively manufactured ABS sheets had about a 17% higher mean average heat release rate (MAHRR). Conversely, the characterization of ULTEM 9085™ sheets revealed the MAHRR of the AM samples were 13% lower than the molded samples. This is attributed to additives in the material used for extrusion AM as well as the build process itself. Effects of geometry were assessed using material cribs, which were composed of layers of rectangular prisms separated by air gaps, with prisms on consecutive layers being orthogonal. Cribs were constructed with three to ten prisms per layer to evaluate the effects of varying the internal material surface area.

Cribs were constructed with three to ten prisms per layer to evaluate the effects of varying the internal material surface area. Below a specific threshold, the burning mass loss rate per unit area of the cribs decreased with an increase in internal material surface area; this agrees with trends predicted using a theoretical model previously developed for wood cribs.ULTEM™ (amorphous1. Introduction1.1 MotivationOriginally founded on the concept of providing a means for prototyping design iterations, additive manufacturing (AM) systems are now being used to fabricate end-use products from both metals and polymers. In addition, a number of companies now produce desktop-scale, inexpensive, extrusion-based AM systems that process polymers such as ABS (acrylonitrile butadiene styrene) and Nylon. 278,000 of such systems were sold in 2015 [1].

278,000 of such systems were sold in 2015 [1]. In addition, the list of processable polymers for industrial-grade AM systems has also expanded in recent years including both low temperature (such as Nylon and ABS) to high temperature polymers such as ULTEM™ and Polyphenylsulfone (PPSF). In addition, systems with larger build volumes have been developed to enable both large scale prototyping and manufacturing large end use components. One such printer can produce objects as large as a car, 10 m by 4.3 m by 3.8 m[2]While additively manufactured polymers are often very similar to those typically used in traditional polymer processing (e.g., extrusion and injection molding), the formulations are often modified to be better suited for the printing process.

For example, Stratasys’ “Digital ABS,” one of a variety of polymers for material jetting platform, has a distinct chemical formulation that2407requires combination of two separate acrylate-based photopolymers that approximate the mechanical properties of ABS, according to its manufacturer[3]. Extrusion AM systems [4] use a thermoplastic form of ABS, but even these formulations are slightly different than those used in injection molding to accommodate specific needs of AM including modified rheology and robustness to humidity.AM technology allows easy formulation of materials with various additives. Most research in AM materials has been focused on improving mechanical properties. However, Lao et al. performed thermal and flammability testing on custom formulations of polyamide 11 nanocomposites, and demonstrated that improved flammability properties can be obtained with specific compositions of fire retardant additives[5].

Their work studied the addition of nanoclays, carbon nanofibers, and nanosilicas into polyamide 11 for production using selective laser sintering.Since AM parts are composed of different formulations of materials, and have different build characteristics (e.g., porosity, poor interlayer adhesion, etc.), than parts made by conventional means, flammability characterization is needed to ensure that AM parts are not significantly different (e.g., more hazardous). As the use of additively manufactured parts increases, and the use of AM becomes more prevalent in end use applications, the fire safety industry will need to assess fire properties accordingly. Such knowledge is needed to assist designers in specifying the use of printed components for end use products. For example, one area of consideration is the large effort by the U.S. Navy to use 3D printed parts on ships[6].

For example, one area of consideration is the large effort by the U.S. Navy to use 3D printed parts on ships[6]. The program, called “Print the Fleet,” aims to significantly save time and resources by allowing replacement parts to be printed on-board (or at a nearby dock) during naval assignment, thus minimizing time invested in resupply. While this may indeed improve the efficiency of ships by allowing longer time on the sea, fire hazards are tightly controlled on ships due to the potential for accidents. Ensuring that AM components will not cause a greater hazard is important to this effort.1.2 ContextThe overall goal of this work is to compare the fire performance of additively manufactured and conventionally produced materials. Heat release rate is often the most important parameter in characterizing the fire hazard of a material or object[7].

Heat release rate is often the most important parameter in characterizing the fire hazard of a material or object[7]. The first objective, therefore, is to explore potential differences between the heat release rate of AM parts and conventionally fabricated (e.g., injection molded) parts. Another objective is to investigate the geometric aspects of flammability. To determine the effect of geometric spacing, the fire performance of material cribs in various configurations will be compared using a theoretical model. Since one benefit to additive manufacturing is the ability to create parts with complex and sparse geometries to save weight, it is important to know to what extent this increases flammability.An overview of the theory of fire safety testing and the model used to correlate the geometric cribs study is shown in Section 2. Tests to perform this characterization use a cone calorimeter, performed according to ASTM E1354[8].

Tests to perform this characterization use a cone calorimeter, performed according to ASTM E1354[8]. The device test method is explained further in Section 3. The heat release rate results and discussion of the material comparison are presented in Section 4. Section 5 contains the conclusion and recommendations based on the results of this work.24082. Theory and Modeling2.1 Fire Safety of PlasticsThe heat release rate of an object in a fire will directly contribute to the hazard. Some materials with fire resistant properties inhibit fire spread, and this corresponds to a low heat release rate. Other materials, often with high carbon, oxygen, and hydrogen compositions, readily burn and cause an even greater hazard once ignited.Fire safety engineers use the heat release rate of objects and components to predict and model fire spread in a room. Buildings and other room-type enclosures (e.g. buses and trains) can then be designed accordingly.

Buildings and other room-type enclosures (e.g. buses and trains) can then be designed accordingly. For example, egress time is sometimes calculated based on how fast a room fire attains full involvement. The heat release rate of objects in a room will dictate how fast the area becomes untenable.Cone calorimeter testing allows characterization of ignition time, heat release rate, smoke release rate, and mass loss of samples exposed to a heat flux. A cone heating element provides a radiative flux to a sample. The combustion gases are extracted and their concentrations analyzed. The depletion of oxygen and the mass flow rate of the combustion gases allows the calculation of heat release rate. Full size objects can be tested in a furniture or other large calorimeter to measure heat release rate and other fire hazard information. However, since those tests are expensive, researchers have developed methods to relate bench-scale results to full scale predictions.

However, since those tests are expensive, researchers have developed methods to relate bench-scale results to full scale predictions. Details on the cone calorimeter used in this study are provided in Section 3.1. The simplest way to extend bench-scale results to larger objects is to use a surface area ratio[9]. Empirical correlations can also be determined using statistical analysis.2.2 Material FlammabilityThe two materials selected for this comparative study are ABS and ULTEM™. ABS (acrylonitrile butadiene styrene) is selected since it may be the most widely used material in AM. Much data on fire properties of ABS are available, and its heat release rate is relatively high compared to many plastics[10][11]. For this reason fire retardants are often added to ABS, and researchers continue to develop new formulations[12][13].

For this reason fire retardants are often added to ABS, and researchers continue to develop new formulations[12][13]. ULTEM™ (amorphous thermoplastic polyetherimide) was developed for high temperature use and good fire resistance, but cone calorimeter data are not readily available. ULTEM™ 9085, the material tested in this study, has been used to demonstrate additively manufactured aircraft seats[14], and the material has passed the Federal Aviation Regulation (FAR) for fire resistance[15].Flammability characterization of the material itself does not provide a full prediction of the hazard. Air gaps and internal geometry are process and design dependent, and they can both promote or inhibit fire spread.

Air gaps and internal geometry are process and design dependent, and they can both promote or inhibit fire spread. This is an important characteristic to measure with parts made via AM, since the process is often used to produce highly complex geometries that feature purposefully designed porosity for lightweighting.24092.2 Examining Geometry Effects on Flammability with CribsIn fire testing, objects with regular air gap spacing and equally dimensioned prisms are termed “cribs.” Much characterization has been done on wooden cribs, most notably by Gross and Block[16][17]. (An example crib is shown in Figure 1.) Researchers have also characterized polymers when burned in a crib configuration[18][19][20]. However, a thorough investigation of most polymers' fire performance when burned in the crib configuration is lacking in the literature. This may be due in part to difficulties with burning polymers in a structured configuration.

This may be due in part to difficulties with burning polymers in a structured configuration. For example, some materials intumesce before and during ignition. In addition, most thermoplastics will melt or collapse before, or during, burning. The second goal of this paper is to investigate the fire performance of the geometry of additively manufactured cribs made from ABS plastic.Figure 1: Crib structure with parameters: square prism (or stick) member has thickness b, spacing s between sticks, number of sticks per layer is n, and number of layers N (the pictured example 4 layers with 7 sticks)The model used in this work is based on Block's theoretical investigation on wood cribs[17]. Block's study on free burning wood crib fires defines a porosity factor to separate two distinct burning regimes, namely the porosity controlled regime and the surface area regime.

Block's study on free burning wood crib fires defines a porosity factor to separate two distinct burning regimes, namely the porosity controlled regime and the surface area regime. In the porosity controlled regime the spacing between the rectangular prisms, as well as the number of layers of prisms, controls the burning rate. In the surface area controlled regime, the spacing does not have an effect, and the burn rate is only dependent on how much of the surface is exposed to the air. Figure 2 contains Block ’s data, which is organized into what he calls the densely and openly packed regimes.2410Figure 2: Block's experimental burning rate data scaled by porosity factor[17]The theoretical model by Block [17] is based on the approximation that the vertical shaft of the crib can be modeled as a porous tube.

After performing a thermodynamics and fluids analysis on this porous tube and wood structure, he found that in the densely pack burning regime the burn rate is a function of the crib height, vent area of the shaft, and cross sectional area of the shaft. These parameters are completely controlled by the square prism (or stick) dimensions, the spacing between the prisms, and the number of layers of prisms. Within a material, all other parameters are constant. These terms are the variables to control the porosity. Above a certain porosity factor, the theory predicts no increase in burn rate when scaled by surface area.Block performed his study comparing mass loss rate (kg/m2·s) for various crib configurations to porosity factor. This work uses a cone calorimeter apparatus to perform the experimentation, and measure the heat release rate, which is the more important parameter in a fire hazard.

This work uses a cone calorimeter apparatus to perform the experimentation, and measure the heat release rate, which is the more important parameter in a fire hazard. Heat release rate can be related to mass loss rate by the heat of combustion of the material.This study seeks to apply this porosity model to AM polymer cribs of similar configuration. Specifically, the burning of ABS cribs is dictated by a porosity controlled regime and a surface area controlled regime. The burning of an object with internal spacing is expected to obey the same physics, whether of wood or polymer construction. Although the pyrolysis properties of wood and polymers differ in terms of char layer development, the general pyrolysis process and fluid flow through the crib is expected to be the same.3. Experimental Methods3.1 Test Method and ApparatusTesting on the plastics was done on a bench-scale cone calorimeter largely according to the ASTM E1354-15a standard[8].

Experimental Methods3.1 Test Method and ApparatusTesting on the plastics was done on a bench-scale cone calorimeter largely according to the ASTM E1354-15a standard[8]. The bench-scale unit was developed to use oxygen consumption calorimetry to acquire heat release data for materials[21]. The device uses the changes in the oxygen concentration (as well as other combustion gases) to determine the heat release of the fire. Cone calorimeters are typically used 1) to provide a comparison of material performance, 2) for2411obtaining material property data, 3) to provide inputs to numerical fire models, and 4) to prove regulatory compliance[21].The apparatus used in this work is shown in Figure 3. The major components of the instrument are identified in the figure. The cone shaped heating element is controlled by a PID controller to provide a constant irradiance to the specimen via an electric resistance heater. A load cell measures the transient mass.

The cone shaped heating element is controlled by a PID controller to provide a constant irradiance to the specimen via an electric resistance heater. A load cell measures the transient mass. The exhaust duct vents the combustion gas mixture from the fire, and part of this stream is sampled for analysis in a paramagnetic gas analyzer. Temperature probes and a pressure transducer on the exhaust stream allow for the calculation of mass flow rate of combustion gases.Figure 3: This photo shows the cone calorimeter used in the ExtReMe Lab at Virginia Tech. The major components are identified in the photo. Crib samples are placed on the sample holder under the cone heater for testing. Tests are performed to ASTM E1354.The testing closely conformed to the ASTM E1354 standard[8]. The apparatus was calibrated each day before testing with a methane burner providing a known 5 kW heat release rate. The samples were placed in a holder with insulation backing.

The apparatus was calibrated each day before testing with a methane burner providing a known 5 kW heat release rate. The samples were placed in a holder with insulation backing. The molded and printed sheet specimens were 4″ × 4″ (100mm × 100 mm) and 0.236″ (6mm) thick. Before the samples were inserted into the apparatus, the cone heater was raised to a steady temperature that provides the desired heat flux. Data were collected for 120 s prior to testing to ensure an accurate pre-ignition state.After the sample was inserted, ignition time was determined by visual observation. This allows for accounting of the delay of the gas analyzer so the heat release rate can be compared to other parameters such as mass loss rate. Flame-out time was recorded, also by visual observation, and the data acquisition system was left on for at least 120 seconds after extinguishment.2412No pilot ignition system was used in this study.

Ignition of the sample was due to the sample reaching its autoignition temperature after being heated with a predetermined irradiance. Common irradiances are 20 kW/m2 for items near a small fire, to over 90 kW/m2 for room fires that are in the post flashover stage[9]. The heat flux value selected for this study was 40 kW/m2. Preliminary testing at 20 kW/m2 revealed inconsistent ignition time.3.2 Material PreparationEvaluating Heat Release RateTo compare printed and molded polymers, three samples of each type were tested. Molded ABS specimens were cut from larger sheet. The printed sheets were made from extruded black and white ABS-M30 in a Stratasys Fortus 400 mc system. Likewise, the ULTEM™ 9805 sheets were printed with the same Fortus printer. The Ultem 9805 molded plaques were obtained from colorxpress.com. All sheets were 6mm thick.Evaluating Geometric Effects on FlammabilityThe geometric flammability study was done using ABS plastic.

All sheets were 6mm thick.Evaluating Geometric Effects on FlammabilityThe geometric flammability study was done using ABS plastic. Crib samples made from molded ABS plastic were manually constructed. Printed crib samples were made using the Fortus described earlier. Support material SR-30 was used in the printing process to scaffold the overhanging features; it was dissolved in a sodium hydroxide solution according to manufacturer’s specifications. The samples were then flushed with tap water to eliminate the solution, and dried before undergoing testing.To ignite the cribs, the top of the crib was exposed to 40 kW/m2 heat flux in the cone calorimeter. The testing done with the plastic cribs was the same as that for the sheet samples. During crib burn testing, mass loss caused the height of the crib to decrease.

The testing done with the plastic cribs was the same as that for the sheet samples. During crib burn testing, mass loss caused the height of the crib to decrease. Since the height of the cone heater was not changed during the test, this increased distance between the cone heater and the sample caused a decrease in the heat flux from the original 40 kW/m2.ABS has a low glass transition temperature of about 100 ºC. At this temperature ABS does not melt into a liquid, but, nevertheless, demonstrates viscous flow. ABS is considered an amorphous polymer, and, as such, it does not have a true melting point. When ABS is burned in a structured configuration, it collapses due to the high temperature. In the experiments performed, all ABS cribs tested did eventually collapse. The cribs with a denser porosity factor maintained their structure longer than the sparsely constructed cribs.

In the experiments performed, all ABS cribs tested did eventually collapse. The cribs with a denser porosity factor maintained their structure longer than the sparsely constructed cribs. To lengthen the time before collapse for lower density cribs, wire supports were constructed (examples of these supports can be seen in Figure 4- b). Collapse was determined by visual inspection to occur from about 20 seconds to over 1 minute.The porosity study was performed with both printed ABS cribs and with cribs made from injection molded plastic. Examples of AM cribs made for the experiments are shown in Figure 4.2413(a)(b)Figure 4: Additively manufactured ABS cribs. (a)The specimen is four layers and six prisms per layer, and the stick thickness is 8 mm. The porosity factor was calculated to be 0.74. (b) The specimen is five layers and two prisms per layer, supported by galvanized steel wire. The stick thickness is 6 mm. The porosity factor was calculated to be 1.75.4.

(b) The specimen is five layers and two prisms per layer, supported by galvanized steel wire. The stick thickness is 6 mm. The porosity factor was calculated to be 1.75.4. Results and Discussion4.1 AM Sheet Polymer ComparisonFigure 5 shows an overall flammability comparison between the heat release rate curves of three additively manufactured polymers, ABS-M30, ULTEM™ 9085, and PPSF/PPSU. All samples were tested according to ASTM-E1354, as described in Section 3.1. It is apparent that the ULTEM™ and PPSF are developed as heat and fire resistant plastics. Evaluating the respective fire hazards reveals that ABS is about 10 times more hazardous in terms of heat release rate. Testing was performed at 40 kW/m2 irradiance, and more testing at different heat flux levels is needed to fully ascertain the relative hazard between these materials. Ignition times between the samples were 20 s for the ABS, 60 s for ULTEM™, and 340 s for the PPSF.

Ignition times between the samples were 20 s for the ABS, 60 s for ULTEM™, and 340 s for the PPSF. The ULTEM™ and PPSF samples displayed intumescent behavior by developing a large char layer as they burned. This both inhibited efficient heat transfer into the material, as well as restricted the upward flow of combustion gases through the cone heater.2414Figure 5: ASTM-E1354 cone calorimeter results for three additively manufactured polymer sheets. Testing was done at 40 kW/m2 irradiance. Flaming for PPSF and ULTEM™ samples did not cease until 1700 s, but steadily decreased during that portion.4.2 AM and Molded Polymer Sheet Heat Release ComparisonFigure 6 shows the comparison of the heat release rate of the AM ABS sheets and conventional molded ABS sheets. The difference between the average values of peak heat release rate is 134 kW/m2.Table 1 shows the comparison of the test averaged heat release rate.

The difference between the average values of peak heat release rate is 134 kW/m2.Table 1 shows the comparison of the test averaged heat release rate. The AM sheets had a higher mean average heat release rate by about 80 kW/m2. Since none of the specimens burned longer than 300 s, the average is for the entire portion of the heat release curve.2415Figure 6: ASTM-E1354 cone calorimeter results for 6mm thick 3D printed and molded ABS sheets at 40 kW/m2 irradiance. The mean peak heat release rate for the 3D printed and molded sheets are represented as the dashed lines in the figure. The average for the 3D printed and molded sheets are 1067 kW/m2 and 933 kW/m2, respectivelyTable 1: Average ABS heat release rates for molded and AM sheets. Data were integrated over 300 s. Units are kW/m2.Test 1 Test 2 Test 3 AverageMolded Sheet 485 470 487 480FDM Sheet 537 523 631 564The resultant ash from each of the molded and printed ABS sheets measured approximately 0.6 g.

Units are kW/m2.Test 1 Test 2 Test 3 AverageMolded Sheet 485 470 487 480FDM Sheet 537 523 631 564The resultant ash from each of the molded and printed ABS sheets measured approximately 0.6 g. Each process produced about the same amount of ash. However, the appearance of each of the resultant ash was markedly different. The ash produced from the molded sheets was whitish in color, while the ash produced from the AM sheets appeared blue. The comparison between the ash samples is shown in Figure 7.This difference between the results for the AM and molded ABS sheets could primarily be due to the distinct chemical formulations. Since the ABS used in the Fortus has been modified for optimal processing via extrusion AM, increased flammability may occur. The density of the molded sheets was greater than the density of the printed sheets by 7%. However, internal density differences between the injection molded and additively manufactured ABS sheet is not believed to have an effect.

However, internal density differences between the injection molded and additively manufactured ABS sheet is not believed to have an effect. Patel et al. performed a numerical study to investigate the effect physical material properties have on cone calorimeter results[22]. Changes in density did not affect peak heat release rate. However, more testing is required to validate this assumption.2416In most cases, the differences observed between the AM and molded material flammability are not significant to warrant extra controls for fire safety reasons.

However, in some applications this trend towards increased flammability of printed ABS may become a more serious consideration, and a better understanding of what causes these materials to exhibit different performance should be studied.(a)(b)Figure 7: ABS sheet combustion ash: (a) AM sheet ash, (b) Injection molded sheet ash.Figure 8: ASTM-E1354 cone calorimeter results for 6mm thick AM and molded ULTEM™ 9085 sheets at 40 kW/m2 irradiance. Only the first 1150 s of data are shown. The mean average heat release rate for the first 1000 s is plotted as a dashed line, with the mean average heat release rate of the 3D printed and molded sheets being 61 kW/m2 and 69 kW/m2, respectively.2417Figure 8 shows the comparison of AM and molded ULTEM™ 9085 sheet heat release rates. The average peak heat release rate is only 2 kW/m2 higher for the printed sheet.

The average peak heat release rate is only 2 kW/m2 higher for the printed sheet. However, the average heat release rate of the molded sheets is 13% higher than the AM sheet over the first 1000 seconds of burning.The ash of the samples is shown in Figure 9. The molded sheets all generally expanded uniformly upward into the cone heater during the test. The AM sheets, however, sometimes seemed to delaminate and expanded erratically. This behavior is likely what caused the lower overall heat release rate in the AM sheets. The figure shows the unburned material which spilled over the side of the sample holder. Although aluminum foil was used to contain the plastic in the sample holder, the aluminum foil was insufficient to prevent this expanded material from flowing off of the sample holder.During the test the material expanded into the cone heater. This intumescent behavior was significant, and caused the material to contact the bottom of the cone heater.

This intumescent behavior was significant, and caused the material to contact the bottom of the cone heater. This inhibited the combustion gases from escaping the top of the cone heater. The large char prevented efficient heat transfer from the cone to the unburned material, and this caused the heat release rate to be low for both AM and molded sheets.(a)(b)Figure 9: ULTEM™ sheet ash from cone calorimeter test (a) AM sheet ash, with unburned material hardened from flowing over the side of the holder, and (b) Molded sheet ash, which was completely contained by the holder and foil wrapping.4.3 Flammability Dependence on GeometryFigure 10 contains the results from the flammability study on the ABS crib tests. The porosity factor from Block's work[17] is plotted against heat release rate. The maximum heat release rate of the portion of the burning before collapse was used, since the airflow properties will change after collapse.

The maximum heat release rate of the portion of the burning before collapse was used, since the airflow properties will change after collapse. CAD images of the cribs used in this study, which demonstrate different porosity factors, are shown in Figure 11.2418Figure 10: ABS crib pre-collapse heat release rate are graphed as a function of Block’s porosity factor. The cribs were tested at 40 kW/m2 irradiance, and both molded and AM produced materials are included in the graph.(a)(b)(c)Figure 11: Cribs used in porosity study, using ABS material; (a) Porosity factor φ = 1.46, (b) φ = 0.81, (c) φ = 0.33The initial tests of the cribs with the larger porosity factors resulted in very short collapse times, and full ignition of the entire crib may not have been realized. This is believed to be the cause of the lower heat release rate values for the molded cribs at porosity factors greater than 1. The wire supports did prolong the structured burn.

This is believed to be the cause of the lower heat release rate values for the molded cribs at porosity factors greater than 1. The wire supports did prolong the structured burn. Most crib configurations were not repeated, but the general trend of the data indicates the applicability of the use of Block's porosity factor to predict two regimes of burning of plastic cribs.2419The constants used in the calculation of the porosity factor were estimated based upon limited data available in the literature. Values of porosity greater than one indicate the burning is in the surface area controlled regime. However, more testing is required to determine exactly how the porosity should be scaled to accomplish a proper quantitative determination of each regime.

However, more testing is required to determine exactly how the porosity should be scaled to accomplish a proper quantitative determination of each regime. Ultimately, the data seem to indicate the presence of two distinct regimes of burning, which demonstrates the importance of designing structures for fire safety performance independent of material characterization.Due to the many factors governing the burning of polymer cribs, the differences between the burning of the AM and molded cribs is not deemed significant. As shown in Figure 6, the initial burning of the AM and molded ABS sheets display similar performance. Only pre-collapse heat release data was used for the porosity model. Most cribs collapsed within 30-50 s, and most of the peak heat release rate values were obtained before 60 s of flaming.

Only pre-collapse heat release data was used for the porosity model. Most cribs collapsed within 30-50 s, and most of the peak heat release rate values were obtained before 60 s of flaming. The sheet data is varied within this region, and differences in material properties therefore do not have as much effect on crib flammability.The data indicate the presence of a porosity controlled regime, and this provides a basis for designing polymer parts for fire performance using geometry. Increasing the resistance to airflow by decreasing the spacing between structural members will reduce the initial heat release rate, and lower the fire hazard of the object. While there are many benefits to creating a sparse structure in terms of material usage, weight reduction, and speed of manufacturing, in certain instances slightly more material being used to block airflow may substantially increase the fire safety of a design.5.

ConclusionAs more designers look to AM as a means of fabricating end use parts, understanding printed parts' fire safety performance becomes imperative. While designers know not to assign material properties of conventionally processed parts to those that are printed (i.e., due to part anisotropy), there remains uncertainty about the flammability of printed materials and their differences with conventionally processed materials.Two tests were performed in this study: AM polymers were compared with their respective molded counterparts, and cribs of various porosities were compared using a theoretical model. Peak heat release rate and average heat release rate comparisons between conventional injection molded and AM parts reveal different behavior between the two processes. The AM produced ABS sheets had an 80 kW/m2 (17%) higher mean average heat release rate than the injection molded pieces.

The AM produced ABS sheets had an 80 kW/m2 (17%) higher mean average heat release rate than the injection molded pieces. The ULTEM™ AM sheet had a 13% lower mean average heat release rate, but its behavior as shown during the test may cause other issues during a fire, as the AM sheets displayed substantial dripping from the holder. The molded sheets did not display this behavior. More testing is needed to fully understand why the AM sheets differ in fire performance than conventionally produced parts.The porosity study revealed two distinct regimes in free burning structures made from ABS plastic. This result indicates that structures may be able to be designed to inhibit fire hazard, even without material controls. Further study is required to fully define the transition point between the porosity controlled regime and the surface area controlled regime.

Flexible File Format for Solid Freeform FabricationStephen 1. Rock Michael 1. WoznyRensselaer Design Research Center Rensselaer Polytechnic Institute Troy, New York 12180AbstractA flexible file format for Solid Freeform Fabrication data is presented which significantly improves on the de-facto industry standard STL format. The new format removes the redundancy present in STL files and can contain topological information. Its specification flexibility allows users to balance storage and processing costs. Since facet boundary models currently provide the greatest common denominator for data exchange between many CAD systems, they are supported by this format. Additionally, representation of CSG primitives is provided, as are capabilities to represent multiple instances of both facet and CSG solids. Format extensibility, without obsoleting existing programs, is made possible by interleaving the format schema with the data.

Format extensibility, without obsoleting existing programs, is made possible by interleaving the format schema with the data. User data can be added to existing entities, or new entities can be created. This allows the addition of NURBS based geometries in the future.1 IntroductionNo standard has been formally adopted for transfering Solid Freeform Fabrication (SFF) data. The STL facet model format specified by 3D Sytems has become the de-facto industry standard in the SFF community [1]. 1.1 STL Format LimitationsSolid objects are represented as an unordered collection of triangular planar facets, and each facet has an outward directed facet normal associated with it.

1.1 STL Format LimitationsSolid objects are represented as an unordered collection of triangular planar facets, and each facet has an outward directed facet normal associated with it. Multiple solids are allowed, although it is unclear from the interface specification whether these are allowed to intersect other solids in the model.Given only the data in an STL file, ensuring a model is physically realizable and is not missing any information (model validity checking) is computationally expensive. Infering model topology from this minimal information is also costly and the are subject to ambiguity caused by numerical imprecision. This suggests the need for an alternative representation which carries greater information content and can reduce model preprocessing demands. STL files also carry a significant amount of redundant information. For each facet, vertex coordinates are specified explicitly. Process attributes are mentioned in the specification but never developed.

For each facet, vertex coordinates are specified explicitly. Process attributes are mentioned in the specification but never developed. Both ASCII and Binary formats exist and it is possible to translate between them. Additional constraints are placed on model data. Ordinate values can never be zero or negative, and all ASCII format keywords must contain only lower-case characters. 1.2 Higher Order Approximationsformat and representation has been The need for an improved SFF mo rec........""·l1 .....:.d [2]. Arguments were made in supp of using surface approximations, termed "precise geometry" by the authors, instead of facet model approximations. Use of a single modeling entity was also recommended. These are valid suggestions; however, to make© 1991 RPI Rensselaer Design Research Center. All Rights Reserved.use of truly precise geometry, information must not be lost by approximation when transfering CAD model data to the SFF process.

All Rights Reserved.use of truly precise geometry, information must not be lost by approximation when transfering CAD model data to the SFF process. Representing models by surface patch approximation will certainly be more compact than planar facet approximation. A surface patch approximation will also experience slower data volume growth as the geometry approxin erroris decreased.1 e sutface approximations are representationally more efficient than their facet other considerations must be analyzed before defining surface patch ations as the new and only SFF modeling entity. Model validity checking will a ing, while having fewer surface patches to intersect, will increase in complexity. Model oximations must be generated, and this too also increase i m lexity. S tessellations are also used in the field of u incurs computer ] and algorithms tailored to SFF are beginning to emerge [4].

S tessellations are also used in the field of u incurs computer ] and algorithms tailored to SFF are beginning to emerge [4]. Code aphi development to support surface patch approximation will be more expensive than that for facet approximation. 1.3 IGES & PDES/STEPpatcost. TriangulCAD system data transfer standardization provides another representational alternative, and some standards exist. The Initial Graphics Exchange Specification (IGES) began as a standard for representing graphical information [5]. It has since evolved to include limited solid model representation capability with CSG primitives supported in Version 4.0 and boundary representation support in Version 5.0. Unfortunately, these representations are verbose [6]. The Standard for Exchange of Product Model Data (STEP) standard is under development and is to represent the complete product model, not just the geometry [7].

The Standard for Exchange of Product Model Data (STEP) standard is under development and is to represent the complete product model, not just the geometry [7]. The generality of this representation is expected to make it verbose with capabilities beyond those needed here. 1.4 2D Slice ContoursThe representational alternatives discussed provide greater information content that the STL Format file. Another alternative has been proposed which provides only slice contour data to SFF processes [8]. Two dimensional slice contours are expected to hold less information than a three dimensional model, and any compression gained by correlation in the third dimension will be lost. However, each slice contour can be provided as a precise representation. This moves the slicing process into the CAD system and eliminates the need to tessellate solid models into approximate facet representations.

This moves the slicing process into the CAD system and eliminates the need to tessellate solid models into approximate facet representations. Unfortunately, this approach limits flexibility at the process controller and will likely increase data transfer volume. New slicing orientations and part nesting arrangements will be difficult to select at the process controller. Slicing in the CAD system effectively moves process control to the processor running the CAD system. Spending design engineering time dealing with SFF process control issues may not be cost effective. This approach also fails to recognize that certain non-CAD data sources, such as reconstructed models from medical imaging data, will likely continue to generate three dimensional facet model data [9].

1.5 Facet ModelsFacet models currently serve as the greatest common denominator between the multitude of CAD systems in existence because many CAD vendors have developed translators capable of generating facet model representations. Facet representations are also very common when transforming scalar fields into three dimensional geometry (medical imaging, remote sensing, etc.)[10, 11]. Facet approximation tolerance thresholds can be decreased to achieve required model precision. While this increases storage demands, the trend of continually decreasing storage cost devalues the consequences.Facet models are not proposed as the optimum solution to the SFF data exchange problem. They are only one alternative but should continue to be supported in future representation schemes because of their greatest common denominator property, both between CAD systems and other sources of geometrical models.

CSG primitives are an2attractive representational addition to a new fonnat because they are well defined. More general standards tend to be verbose. This suggests a data exchange format tailored to SFF applications is appropriate. In. this paper, we p:esent a flexible, exte~sibl~ mode~ format which supports both facet solIds and eSG solIds. It supports multiple InstancIng of a single solid object definitions as well as a number of SFF specific entities. Topological information is provided for facet models to increase validity checking and processing efficiency.2 RPI Format Design ConsiderationsA number of design considerations for an improved SFF format, termed RPI format, are presented in this section. The RPI format must be derivable from currently accepted STL format data. A radical change which fails to maintain such com ity will gain poor acceptance. The RPI format must provide an extensible framework for evolution.

A radical change which fails to maintain such com ity will gain poor acceptance. The RPI format must provide an extensible framework for evolution. This includes provisions for adding new geometric entities, adding data to existing entities, and adding non-geometric process data. The new format must also allow flexibility, with respect to both format conformance and the amount of data stored. It is also important to eliminate the redundancy present in STL format files. 2.1 Upward Compatibilityformat is possible. The RPI format format to is capable of representing facet solids, but it includes additional information about facet topology. Topological information is maintained by representing each facet solid entity with indexed lists of vertices, edges, and faces. Instead of explicitly specifying the vertex coordinates for each facet, a facet can reference them by index number. This contributes to the goal of overall redundant information reduction.

Instead of explicitly specifying the vertex coordinates for each facet, a facet can reference them by index number. This contributes to the goal of overall redundant information reduction. 2.2 Topological Information model,Upward compatibility frominformation can be inferred at some computational expense. This is also true for the RPI format and is the mechanism by data which data volume flexibility is provided. Many possible combinations of t vertex can be maintained for facet solids in an RPI format file. In the simplest list and face list are provided. The vertex list is referenced directly by dex values in the face list. When normal data is included, this is informationally equivalent to the STL format. It is not even necessary to specify facet normals as these can be computed based , a on vertex ordering assuming the right hand rule is followed. In its most complete st, and face list. The ed facet model can be represented by a vertex list, edge can ne each edge.

In its most complete st, and face list. The ed facet model can be represented by a vertex list, edge can ne each edge. The face list can reference the two vertices and two faces which d faces. Normals can also be reference each faces vertices, edges, and three neigh emes is specified for each face. Any range of to n between will be logical· permissible. When less t required to generate it. is flexibility llows storage and proc ssing costs to be appropriately balanced. 2.3 Conformance and ExtensibilityGiven antopological,specifiation is provided, moreformat conformance flexibility and format extensibilit coupled objectives. Both are achieved by imbedding the format schema wit The schema defines the data which represents each e names and defines the order in which they occur, but it also each of these variables. This allows unrecogni varia irable t renderin the remaini data, as id ould in where omitting unrecognized data is very useful.

This allows unrecogni varia irable t renderin the remaini data, as id ould in where omitting unrecognized data is very useful. For example, the Phong Shading utilizes vertex normal information which could be s re using the RPI format [12]. This same file should be readable by an SFF processor which is completely unaware of the existence of, or variable names used for, vertex normal data in the file. By associating a data type with each variable name in the file, data corresponding toProvi. It not 0efiness the d be skta unr mble. It is not a ways error; however,th3unrecognized keywords can be read and discarded. This mechanism also allows entire unrecognized entities to be skipped; a feature of the public domain Interchange File Format (IFF) specification [13].

This mechanism also allows entire unrecognized entities to be skipped; a feature of the public domain Interchange File Format (IFF) specification [13]. This allows the format to evolve without rendering existing interpreters useless with each new format addition.Imbedding format schema information in the data also allows the order in which Instead of enforcing specific data record ordering, the data is prese variable names ound in each schema definition determine the order in which data is flexibility also exists at the entity level. Most model entities can appear presented. S it is the interpreter's job to derive the necessary information. This in any orde provides conformance flexibility while placing a greater burden on the RPI format interpreter.to be altered.entities can be added to the format and new variables can be added to exiting ,ho ver can neither be added nor modified while preserving the ta entities.

If an interpreter reads an unknown s' b it can successfully c processing the data because it knows what to read and v If an unknown data type is used, the interpreter will not know how much discard. information to read and is unable to continue processing. Consequently, data type selection during format specification is critical to ensure the format's successful evolution. The data types provided closely resemble those available in the C programming language [14]. Integers, single and double precision floating point values, strings of characters, and boolean operations are supported. Additionally, a binary type is defined which supports binary values of any length. 2.4 CSG Primitivesar compa 1 ilityed into the format.In addition to facet solids, the RPI format supports cuboid, cylinder, cone, sphere, and tori CSG primitives. These are common in mechanical design and one study found that more than 90% of an industrial sampling of parts could be constructed from them [15].

These are common in mechanical design and one study found that more than 90% of an industrial sampling of parts could be constructed from them [15]. The primitive representation is also much more efficient than the corresponding facet approximation. Transformation entities are provided so that actual solid coordinates do not have to be individually repositioned. This is useful when orienting CSG primitives or when attempting to nest larger collections of solids. The same solid or collection of solids can be instanced multiple times without replicating the data. This serves to reduce data redundancy and provide flexibility. Finally, non-geometrical entities are provided. These allow process data to be maintained within the RPI format file. Included are entities which represent the process type, scanning methodology, and provide notes for the machine operator.The benefits encompassed in the RPI format are not without cost.

Included are entities which represent the process type, scanning methodology, and provide notes for the machine operator.The benefits encompassed in the RPI format are not without cost. Developing an interpreter which processes a format as flexible and extensible as the one described is non trivial. However, from a global perspective on effort, it seems complicating interpretation for a few SFF producers is worth the benefit of simplifying format conformance, and increasing storage and processing efficiency, for many times that number of end-users.3 The RPI FormatAn abridged specification of the RPI format is provided to illustrate the concepts previously discussed. The format was developed in ASCII to facilitate cross-platform data exchange and debugging, but attention has been given throughout format design to ensure a reasonable binary representation mapping is possible.

An RPI format file is composed of a collection of entities, each of which internally defines the data it contains. Each entity conforms to the syntax defined by the syntax diagram shown in Figure 3.1[16].4SchemaType Definition3.1 RPI FormatEach entity is composed termination symbol, and the "'...........""" .....""'..... rt ..... ..,. records which are made by the Schema. EachanVariable types inclludednt:ege:r Ifloat (DBL), string (BITn) where n is a nrt.c·~'1-~"'T&JA ~1t"'l,T"'lT'::'''' newline, etc.), so of the standard escape s variable type, or integer or the"?" """"..........,....""'" precedes. An integer repeat count which v causes the data it defines to subsequent fiel variable names to be rt.Tn1f"hl.rt that variable names schema termination symbol, syntax diagram for the definition. C style comments (/* quoted strings where 3.1 Entityto beis notwill be _""fl'"C'''Eions tions existtransforma.

DeVariables:INDEX VI, V2 (INT) - defines vertex index numbers for the two vertices which define an edge. Fl, F2 (INT) - defines face index numbers for faces meeting at the edge.) - defines the index number for each edge.FACES This defines the faces used by a facet solid representation. Defined Variables:INDEX (INT) - defines the index number for each face. VI, V2, V3 (INT) - defines vertex index numbers for the three vertices which define a face. El, E2, E3 (INT) - defines edge index numbers for the three edges which define a face. F12, F23, F31 (INT) - defines face index numbers for faces adjacent to the face being defined. The numbering convention indicates the which edge a neighboring face is sharing. NX, NY, NZ (DBL) - define the normal direction vector for the face.CUBOID, CYLINDER, CONE, SPHERE, & TORI These entities define the five eSG primitive types supported. Dermed Variables:INDEX (INT) - defines the index number for each primitive.

Defined Variables: FILE (STR) - defines the filename, including path, for inclusion.3.2 A Facet Model ExampleAn example of facet model definition using theFormat is provided. The dataprovided is an excerpt from the 7160 facet model referenced in Figure 4.1./* The following file is a .RPI Flexible Format Facet File */ PROCESSDATA 1 PART_NAME STR * tooth.

features should not be ignored when apI)lic:aticm needed to repeat count capability w ld be very useful. 4 RutI lzengfor theformat.s ed which allows RPI format files to be tationally expensive task because model model facets. All work has been done wn should remain valid when binary formats are se the ASCII and Binary representations do not differ significantly in Format files to corresponding RPI n. Figure 4.1 contrasts the file size ofbeen is . the unorderedin at files. ComparisonslIes.KEY: ~ STL FormatD Minimal RPI FormatI CD e. C1> .1:::1 C/) ~ L.CFull RPI Format1m STL Equivalent RPI FormatCompressed Files4045106161192Test Part Facet Count66547160Test Part Facet CountFigure 4.1 - STL Format/RPI Format File Size ComparisonsThe Full RPI Format file contains the full range of topological connectivity information needed to make model slicing more efficient with the topological slicing approach of [17].

The experimental topological slicer implementation interprets a sub-set, supporting only single facet solids, of the RPI format specification previously defined to avoid the costly step of inferring model topology from an file. The STL Equivalent RPI Format file contains information content equivalent to that of an STL file, and the Minimal RPI Format file contains the minimal information necessary for model representation. Only the Full RPI Format file is suitable for active use; however, since it can be generated at some computational expense from one of the other formats, this allows8a balance between storage and processing costs. The Minimal RPI Fonnat file should prove useful for archival storage.The UNIX Compress utility was used to compress each of the files using adaptive Lempel-Ziv coding [18]. The amount of compression provides some indication of how much redundancy is present for each of the formats.

The amount of compression provides some indication of how much redundancy is present for each of the formats. Notice that while the Full RPI Format file averages approximately 50% the size of the corresponding STL Fonnat file, the compressed STL file is typically marginally smaller than the compressed Full RPI Format file. This highlights the redundancy present in STL Fonnat files. Also note that the compressed STL Equivalent RPI Fonnat file is smaller than the compressed STL file in all experimental cases shown. This suggests it may be suitable for archival purposes.5 Conclusion5.1 ConclusionsThe RPI Fonnat offers a number of features unavailable in the STL Format. The fonnat can represent CSG primitive models as well as facet models. Both can be operated on by the boolean union, intersection, and difference operators. Provisions for solid ocess translation and multiple instancing are also provided.

Both can be operated on by the boolean union, intersection, and difference operators. Provisions for solid ocess translation and multiple instancing are also provided. Process parameters, such as type, scan method, material, and even machine operator instructions, can be inclu .n the file. Facet models are more efficiently represented as redundancy is reduced. The flexible format definition allows storage and processing costs to be balanced.While these features are useful, perhaps the largest benefit of this fonnat is its extensibility. It provides for backward compatibility and allows users to include additional non-SFF data, or new SFF data, in a file. Emphasis was placed on RPI Fonnat design considerations, not on specification of the defined entities, because this paper does not attempt to define a new standard. A sub-set of the format specified has been used to our benefit; however, it is by no means complete.

A sub-set of the format specified has been used to our benefit; however, it is by no means complete. The goal of this work is to provide an extensible format which can evolve with new representational enhancements. This should lessen the number of distinct file fonnats in use and eliminate the translators necessary to convert between such formats. 5.2 Future WorkIn the near tenn, a binary fonnat equivalent to the ASCII RPI fonnat presented must be developed. Model precision requirements should be investigated to detennine what representational precision is necessary in a binary fonnat. This is less of an issue in ASCII fonnat because any precision can be represented. It would also be useful to develop a method for creating new data types which do not disable backward compatibility.Research must be conducted to identify a surface patch suitable for solid approximation.

It is also important to develop basic algorithms which can approximate CAD system entities with these surface patches. This resurrects the issue of primitive level CAD system interfacing and the cost associated with developin uch an interface. As the PDES/STEP standard is developed, it should be evaluated as the potential primary representation for SFF data transfer. It may be appropriate to allow a varlet of entry e process points to SFF processes; from high level CAD data thr controller.

It may be appropriate to allow a varlet of entry e process points to SFF processes; from high level CAD data thr controller. Advancements in controller technology may also influence the direction of future SFF model fonnats.scan vector input toAcknowledgementsThis research was supported by NSF Grant DDM-8914212 as a subcontract through the University of Texas Solid Fre ts of the Rensselaer Design Research Center ( Any opinions, findings, conclusions, or recommendations expressed in this publication are those of the authors and do not necessarily reflect the views of the National Science Foundation or any of the industrial sponsors..m, and 0 ciates Pr9We would like to thank Dick Aubin, Pratt & Whitney (a division of United s), for providing a number of industrial STL models. Research by James CT scan data ral Electric Corporate Research velopment.

Research by James CT scan data ral Electric Corporate Research velopment. A James Miller for all the valuable comments and ideasTechnoloerated the facet model of the tooth nerve cavity10REFERENCES1. "Stereolithography Interface Specification," 3D Systems, Inc., June 1988.2. James Darrah and Martin Wielgus, "A New CAD Model Format for SFF Machines?," Soild Freeform Fabrication Symposium Proceedings, University of Texas at Austin, Aug. 1990.3. Reed D. Clay and Henry P. Moreton, "Efficient Adaptive Subdivision of Bezier Surfaces," Eurographics 188, Elsevier Science Publishers B.V., 1988.4. X. Sheng and U. Tucholke, "On Triangulating Surface Model for SLA," Second International Conference on Rapid Prototyping, Dayton, OH, 1991.5. "The Initial Graphics Exchange Specification (IGES) Version 5.0," National Institute of Standards and Technology, Gaithersburg, MD, 1990.6. P.R. Wilson, LD. Faux, M.C. Ostrowski, K.G.

"The Initial Graphics Exchange Specification (IGES) Version 5.0," National Institute of Standards and Technology, Gaithersburg, MD, 1990.6. P.R. Wilson, LD. Faux, M.C. Ostrowski, K.G. Pasquill, "Interfaces for Data Transfer Between Solid Modeling Systems," IEEE Computer Graphics and Applications, Jan. 19857. "Product Data Exchange Specification (PDES) Testing Draft," National Institute of Standards and Technology, Gaithersburg, MD, 1988.8. Richard J. Donahue and Robert S. Turner, "CAD Modeling and Alternative Methods of Information Transfer for Rapid Prototyping Systems," Second International Conference on Rapid Prototyping, Dayton, OH, 1991.9. James V. Miller, "On GDM's: Geometrically Deformed Models for the Extraction of Closed Shapes from Volume Data," M.S. Thesis, Rensselaer Polytechnic Institute, Dec. 1990.10. William E. Lorensen and Harvey E. Cline, "Marching Cubes: A High Resolution 3D Surface Construction Algorithm," Computer Graphics, Vol. 21No. 4, July 1987.11.

1990.10. William E. Lorensen and Harvey E. Cline, "Marching Cubes: A High Resolution 3D Surface Construction Algorithm," Computer Graphics, Vol. 21No. 4, July 1987.11. Lori Scarlatos and Theo Pavlidis, "Hierarchical Triangulation Using Terrain Features," Visualization 90 Conf. Proc., Oct. 1990.12. J.D. Foley, A. vanDam, S.K. Feiner, J.F. Hughes, Computer Graphics Principles and Practice, second ed., Addison-Wesley Publishing Co., 1990.13. Reference, pp. 1-5 - 1-24, 1985.'''EA IFF 85' Standard for Interchange Format Files," Commodore-Amiga Technical14. Brian W. Kernighan and Dennis M. Ritchie, The C Programming lAnguage, Prentice Hall, Englewood Cliffs, NJ, 1988.15. MJ. Pratt, "Solid Modeling and the Interface Between Design and Manufacture," IEEE Computer Graphics and Applications, Jul. 1984.1116. Derick Wood, Theory of Computation, John Wiley & Sons, Inc., 1987.17. Ste hen J. Rock and Michael 1. Wozny, "Utilizing Topolo ical Information to Incr Sy can18.

FLEXIBLE RAPID PROTOTYPING CELLCheukfung Lai and Ian GibsonCenter for Advance Product Development Technologies Department of Mechanical Engineering The University of Hong Kong Pokfulam Road, Hong KongAbstractRapid prototyping systems have demonstrated high flexibility in terms of creating complex geometry parts. Improvements in accuracy and material properties enable rapid prototyping to become a widely used and important part of the product development process. Further applications and further improved performance can be achieved by combining rapid prototyping machines with conventional machine tools to form flexible manufacturing cells. One way to form such a cell is to integrate with industrial robotics. This paper will describe work carried out using an ABB IRB 1400 industrial robot in conjunction with a DTM Sinterstation 2000 to form a Flexible Rapid Prototyping Cell.Keywords: Rapid Prototyping (RP), Selective Laser Sintering (SLS), Surface Finishing, Robot.1.

INTRODUCTION Industrial robots play an important role in automated manufacturing systems. In such systems, materials processing machines are linked together by industrial robots and automated guided vehicles. It is logical to consider Rapid Prototyping (RP) as part of a longer process chain, which makes use of RP’s speed and ability to create complex geometry. Robots are capable of integrating with conventional RP machines to produce flexible manufacturing cells [1]. Such cells could be regarded as Flexible Rapid Prototyping Cells (FRPCs). The purpose of an FRPC is to achieve further automation in prototype manufacturing. Rapid prototyping machines are highly flexible in terms of creating complex geometry. In many cases prototypes built by RP machines have to go through a series of post-processes before they can be delivered to the customers.

In many cases prototypes built by RP machines have to go through a series of post-processes before they can be delivered to the customers. In the case of DTM’s Selective Laser Sintering (SLS) process, the most important post-processes are prototype removal, prototype handling, waste material handling and surface finishing of prototype. The handling of finished parts as well as unused materials have to be carried out manually. Due to variations in the SLS process and geometry related issues, tolerances within the prototype will vary and human errors in post-processing may also result in poor overall accuracy. Furthermore, the amount of manual effort spent on post- processing restricts the usefulness of RP technology. Because of the wide range of purposes for RP parts and the geometrical complexity, most RP post-processing must be performed manually.Rapid prototyping machines are not designed to work automatically with other specified working platforms.

However, re-engineering a rapid prototyping machine so that it was capable of working with a robot would be a relatively trivial task in many cases. Some robot configurations275are specifically designed to mimic human arms (revolute coordinate systems) and as such can carry out many of the complex handling tasks normally restricted to manual operations. This paper will go on to discuss modifications that can incorporated to solve some of the problems in the SLS process. Previously manual tasks can be facilitated by the integration of a robot arm on SLS RP technology. Experiments were focus on a DTM Sinterstation 2000 and ABB IRB 1400 industrial robot.2. BRIEF REVIEW ON SLS POST-PROCESSES After a prototype is built in the SLS machine, it has to go through a series of post-processes, which are usually carried out manually. First the powder cake must be transferred from the SLS machine to the Rough Break Out (RBO) station.

First the powder cake must be transferred from the SLS machine to the Rough Break Out (RBO) station. The RBO station assists with sorting out parts from loose powder and collecting unused powder. Prototypes are then moved to an abrasive- blasting machine where powder that sticks on the prototype surface is removed. Powder trapped in tight corners and holes is not easily removed by this method and manual cleaning of the part is often required. Coating and infiltration can help improve the surface quality, which is normally achieved by dipping the prototype into a polymer resin solution and then drying it. Finishing the prototype to the required roughness is usually achieved by using abrasive paper. Finally primers, paints and other surface coatings are applied if required.3.

Finishing the prototype to the required roughness is usually achieved by using abrasive paper. Finally primers, paints and other surface coatings are applied if required.3. HANDLING OF PARTS To obtain parts from the SLS machines automatically would not be a difficult task, if all that is required is simple transport of the whole powder cake to the RBO station and separation of the prototypes and powder. However if it were necessary to carry out other automated tasks, then it would be useful to have some method for maintaining registration of position and orientation of the prototype. Using a support system can solve this problem. Many existing RP machines use supports to fix parts to the build platform, maintaining position and orientation relative to the platform. Examples include Stereolithography (SLA) and Fused Deposition Modeling (FDM). SLS does not normally require supports because unsintered powder surrounds the prototype, giving it an advantage over other processes.

SLS does not normally require supports because unsintered powder surrounds the prototype, giving it an advantage over other processes. However when sorting out SLS prototypes from the powder cake, powder must fall away and parts will shift in relative position. Using a robot arm to catch the prototype at this time would require sophisticated detecting devices and control units. Even so, the robot may be able to detect and clamp the part, but may not be able to maintain registration on the position and orientation. 3.1 Support Structures The role of the support is to link the object with the part piston. As a result, the relative position and orientation between the part and the part piston is always known. The primary function of SLA and FDM supports is to maintain the position during the build time and prevent the collapse of overhanging structures. Since the purpose of the SLS supports is different, they will have different design concepts.

Since the purpose of the SLS supports is different, they will have different design concepts. The supports used in SLS for the FRPC would have the following requirements:1. 2.Maintain the part’s relative position and orientation to the part piston. Provide a means for the robot arm to clamp the part accurately.2763.2 Fixture In this project, a fixture system was designed to fulfill the above requirements. Fig3.1a shows the components of the fixture system. The fixture system consists of a platform, a handle for the robot, and a base. The base is bolted to the Sinterstation part piston. The platform and the handle are fixed together to form a single part, and there is no relative movement between the two. There are pinholes under the handle, which mate with the pins above the base. Once the handle is fitted, only upward movement is allowed.Fig 3.1a, CAD drawing of the fixture system showing the platform (top piece), handle (middle), and the Pin is (bottom).

Once the handle is fitted, only upward movement is allowed.Fig 3.1a, CAD drawing of the fixture system showing the platform (top piece), handle (middle), and the Pin is (bottom). Fig 3.1b, installed inside the Sinterstation 2000.the fixtureFig 3.1aFig 3.1bBefore building the parts, the fixture system is installed into the machine. The machine startup process is mostly the same as usual, except the level of the part piston is adjusted until the platform surface of the fixture system is level with the build plane. The inclination and the flatness of the platform should be carefully adjusted since any part of the platform higher than the lowest level of the roller will result in build failure.

The inclination and the flatness of the platform should be carefully adjusted since any part of the platform higher than the lowest level of the roller will result in build failure. However no part of the platform can be too low, because if this happens the first layer of powder above the platform will be too thick to penetrate by laser and the part will not adhere to the platform.Experiments have been carried out to prove that it is possible to start sintering on such a platform. Figs 3.2a and 3.2b show blocks attached to a platform. In order to achieve this, the piston level has to be adjusted precisely. Furthermore different laser powers and scanning speeds have been tested. For the DTM Sinterstation 2000 used, laser power was set to 14 (about 9W), having normal scanning speed, when Duraform is used.

For the DTM Sinterstation 2000 used, laser power was set to 14 (about 9W), having normal scanning speed, when Duraform is used. This is higher than the normal laser power used for Duraform, which is around 7W.Sintered materialPlatformFig 3.2aFig 3.2bThe function of sintering the first layer of powder above the platform is to build a support structure between the parts and the platform. This structure provides a rigid linkage between the part and base. Fig 3.3a to 3.3e show the sequence of operations.277Fig3.3aFig3.3bFig3.3cFig3.3dSetup the fixture as shown in Fig 3.3a, and then immerse the system in the powder as shown in Fig 3.3b. Control the layer thickness above the platform. Start building in the chamber. Finally the part is built with the support attached to the platform (Fig 3.3e).Fig3.3e4. REMOVAL OF WASTE MATERIAL Two strategies are applied to remove the material from parts.

Finally the part is built with the support attached to the platform (Fig 3.3e).Fig3.3e4. REMOVAL OF WASTE MATERIAL Two strategies are applied to remove the material from parts. The first approach is used to remove large amounts of unused powder for recycling. Installing a built-in vacuum cleaning system into the SLS machine is an obvious solution. The only problem would be to redesign the build chamber to reduce the effect of contaminating structures inside the chamber. This process aims to remove up to 80% of the unused powder. The following section shows an example of how to redesign the part piston, based on the concept of pipe installation. Fig 4.1 shows that the pipes are installed to surround the part cylinder. These pipes are connected to filters and a suction device. The system is constructed to suck the powder whilst the part piston is rising. Fig 4.2 shows the operation of the system.

These pipes are connected to filters and a suction device. The system is constructed to suck the powder whilst the part piston is rising. Fig 4.2 shows the operation of the system. By using this built-in design it will be very difficult to upgrade the existing equipment, but the concept can also be realised by using an external device. Fig 4.3 shows the structure of the external device. Because several kinds of powder can be used in the SLS machine, changing the external device eliminates the problem caused by contamination of powder inside the pipes.Fig4.3 shows a transparent shell with pipes connected as an external device. The working principle is similar to the one shown in Fig4.2Fig4.1Fig4.2Fig4.3Using this system, it may be difficult to remove the powder close to the parts’ surface. This will be done after the parts and the fixture have been taken out of the Sinterstation.

This will be done after the parts and the fixture have been taken out of the Sinterstation. Development of a brushing machine to clean the parts in all directions is one possible solution for further removal of powder. The idea of using sand blasting is also possible by installing a nozzle onto the robot arm. A further possibility would be to use centrifugal force by rotating the part at high speed, so that the powder on the parts’ surface will eventually detach.2785. REMOVAL OF PARTS FROM THE MACHINE There are occasions when it is not necessary to use a part fixture within the Sinterstation. If we simply want to collect the parts, and do not care about their position and orientation after they have been taken out, then the support and fixture system mentioned before is not required. However by doing this, registration of position and orientation will be lost.

However by doing this, registration of position and orientation will be lost. So, as far as possible, a fixture should be included as standard equipment of an FRPC and a support must be built unless the geometry of the part is too difficult to have supports under it. Recall the procedure used so far. After using the vacuum cleaner to extract most of the powder for recycling, the fixture and the parts would be taken out by using a suitably equipped robot arm or other transporting device. Another advantage of using a fixture is therefore the elimination of the need to design grippers for clamping different kinds of parts.The parts and the fixture are taken out and placed on another working platform. In this working area, cleaning and finishing of parts will be carried out. The fixture would be used to register the position of the part in the work area.6. PARTS FINISHING6.1 The part accuracy There are four major elements that may affect the part accuracy.

The fixture would be used to register the position of the part in the work area.6. PARTS FINISHING6.1 The part accuracy There are four major elements that may affect the part accuracy. The CAD model approximation, distortion caused during building, stair-step effect on curved surfaces, and surface roughness resulting from the process used [2].CAD model approximation errors are differences between the desired model geometry and model geometry obtained by the RP machine. The popular STL file format uses triangular facets to approximate the actual model. When dealing with straight lines and planes, the error is small, but if STL is used to approximate curved or freeform surfaces the tessellation process can generate larger errors. Increasing the mesh density can minimize this problem, but the file size will become larger since the number of triangles required representing the surface is increased [3].Distortions are changes of geometry during or after the building process.

Curling, Glazing and Bonus-Z are common problems in SLS machines [4]. These problems are reducing as a result of improvement in properties of new materials.Stair-step is a problem faced by most RP machines. Since nearly all RP machines use the concept of layering, if the layer thickness is too high then the error is large and the resolution decreases. If the layer thickness is decreased, the errors can be reduced but the build time increases. Sometimes reorientation of building direction can help reduce the stepping problem.6.2 Conventional parts finishing Most SLS parts will go through a manual finishing process aimed at improving the surface quality. Abrasive papers and other additives are applied in this case. Finishing of parts can be achieved by using additive techniques and subtractive techniques. Additive techniques may include the application of coatings, protective layers or surface strengtheners.

Finishing of parts can be achieved by using additive techniques and subtractive techniques. Additive techniques may include the application of coatings, protective layers or surface strengtheners. These coatings or279layers may be removed after all the processes have finished or just left on the surface. Subtractive techniques may include milling, turning, drilling, polishing, blasting and grinding. All these process may cause a loss of material. Also it is worth mentioning that both additive and subtractive methods can be applied in a selective area or over the entire surface.6.3 Robot finishing Most SLS prototypes have a need of surface polishing and finishing in post-processing to improve the surface quality [5]. The geometric sensitivity of robots provides sufficient potential for performing finishing tasks.

The geometric sensitivity of robots provides sufficient potential for performing finishing tasks. When compared with CNC, robot arms usually have larger work envelopes and more degrees of freedom, which enable them to tackle tasks that require complex geometrical manipulation. The tasks here are similar to those used in robot sculpting and finishing [6]. Besides milling and grinding, the robot arm can also act as a device for applying coating in selected regions.Experiments have been carried out to investigate the effects of coating and the robot finishing process on surface roughness of the prototypes. Prototypes made of nylon composite have been used. The procedure was to apply coating on the prototype’s surface, and then tools held by the robot arm finished the surface. The coating material used was a cyanoacrylate based adhesive. The tools used were modified grinding tools with spring-loaded mounting. Coating material infiltrated into the porous structure of the surface.

The tools used were modified grinding tools with spring-loaded mounting. Coating material infiltrated into the porous structure of the surface. Voids on the surface were blocked and the surface hardness increased. Durometer Type D was used to measure the hardness of the surface and it was recorded that after applying the coating the hardness increased from 62 to 75. These characteristics help achieve lower surface roughness after finishing since hard and brittle surfaces can attain better surface quality after finishing than soft and tough ones.It should be noted that if voids appear on the surface, the roughness will increase. Fig 6.1 shows the cross-section of a specimen finished without applying coating. Voids appear on the surface. Fig 6.2 shows the cross-section of a specimen coated with adhesive. Voids on the surface are occupied by the adhesive.

Voids appear on the surface. Fig 6.2 shows the cross-section of a specimen coated with adhesive. Voids on the surface are occupied by the adhesive. It can be seen that application of the adhesive reduces the number of voids appearing at the surface after finishing.Fig 6.1Fig6.2After a model is designed, two kinds of files are generated for different purposes. These files are STL and Cutter Location Source File (CLSF) files. RP machines read STL files for building the prototype. CLSF files record tool paths generated by CAM software. These contain parameters like tool positions and orientations. Tool position is described by a tool center point in the280Cartesian coordinate system, while tool orientation is described by vector notation. For robot finishing, a CLSF file cannot be used directly. The robot controller has its own programming language named RAPID [7].

For robot finishing, a CLSF file cannot be used directly. The robot controller has its own programming language named RAPID [7]. In RAPID, the user defines a user coordinate system and an object coordinate system to coincide with the coordinates used in CLSF file. The relative position and orientation is determined by calibration of the workpiece and the tool coordinate system is determined by calibration of the tool. RAPID uses Euler angles to describe the orientation of the tool and uses a quaternion to describe the orientation of the workpiece.The tool path generated is translated from CLSF file to a RAPID (PRG) file. The robot has 6 degrees of freedom. 5 degrees of freedom are sufficient to determine the position and attitude of the tool [8], since the tool is rotational and its axis of rotation is along the 6th axis.

5 degrees of freedom are sufficient to determine the position and attitude of the tool [8], since the tool is rotational and its axis of rotation is along the 6th axis. To finish curved surfaces with the tool axis normal them, a 5-axis robot finishing technique is required.Fig 6.3 shows the layout of the robot finishing platform. A prototype is fixed on a fixture, which is located within the working area of the robot. The robot carries the tool, which is spring-loaded. Fig 6.4 shows the details of a spring-loaded tool. This kind of tool restricts the pressure applied to the surface. The pressure can be adjusted by changing the spring. In experiment the pressure used was between 20 and 30 kPa. Use of a spring-loaded tool also enabled better contact between the tool and the surface. This is important since it is assumed the prototype built by RP machine will have dimensional errors due to different kinds of distortions.

This is important since it is assumed the prototype built by RP machine will have dimensional errors due to different kinds of distortions. If a rigid tool is used, the tool tip is difficult to trace over the surface. Interference and clearance may occur during the process.Fig6.3Fig6.47. CONCLUSIONS A method for integrating a DTM Sinterstation 2000 with an industrial robot (ABB IRB1400) has been described. It has been found that this combination is feasible to realize a Flexible Rapid Prototyping Cell (FRPC), because most of the technologies required have already been developed. The prototype can be built using a fixture platform in order to maintain the prototype’s position and orientation. Finishing of surfaces can be automated by industrial robot. The conditions for using a fixture within the Sinterstation have been studied.

Finishing of surfaces can be automated by industrial robot. The conditions for using a fixture within the Sinterstation have been studied. This includes the design of the fixture, the process environment (part bed temperature and laser power) required and the control of part piston height.5-axis robot finishing techniques have been studied. This includes the orientation transform function between the CLSF file and robot PRG file, because CLSF uses tool axis vector notation while PRG uses Euler angle system. The calibration of tool and workpiece has been studied as281well. Furthermore the use of compliant tools has been studied.The use coating to improve prototype surface quality has been studied. Application of cyanoacrylate based adhesive on the prototype surface is followed by abrasive finishing.

Application of cyanoacrylate based adhesive on the prototype surface is followed by abrasive finishing. Roughness is decreased when compare with abrasive finishing alone.The combination of the SLS RP technology with robot handling and sculpting appears to be a technical problem that can be achieved. Further study will attempt to establish how much time and manual effort can be saved using this approach.REFERENCES 1. Ian Gibson, “A discussion on the concept of a flexible rapid prototyping cell”, RapidPrototyping Journal, Vol. 2 No. 2, 1996, pp 32-38.2. Dongping Shi, Ian Gibson, “Surface Finishing of Selective Laser Sintering Parts with Robot”, Solid Freeform Fabrication Symposium 1998, pp 27-36.3. Georges M. Fadel and Chuck Kirschman, “Accuracy issues in CAD to RP translations”, Rapid Prototyping Journal, Vol. 2 No. 2, 1996, pp 4-17.4. DTM, The Sinterstation System, Guide to Materials: Duraform Polyamide, DCN: 8001- 10014, December 1997.5.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFLEXURAL BEHAVIOR OF FDM PARTS: EXPERIMENTAL, ANALYTICAL AND NUMERICAL STUDYMadhukar Somireddy*1, Diego A. de Moraes*, and Aleksander Czekanski* *Department of Mechanical Engineering, York University, Toronto, ON, M3J 1L4, Canada 1Corresponding author: madhukar@yorku.caAbstractFused deposition modelling (FDM) processed parts behave as composite laminate structures. Therefore, mechanics of composite laminates can be adopted for the characterization of mechanical behavior of the printed parts. In this study, the flexural properties of the 3D printed laminates are investigated experimentally, analytically and numerically. Each layer of the printed specimens is treated as an orthotropic material.

Each layer of the printed specimens is treated as an orthotropic material. The elastic moduli of a lamina are calculated by considering the mesostructure of the printed laminate in finite element simulation of tensile testing. These elastic moduli are employed in a constitutive matrix for the calculation of flexural stiffness of the laminate using classical laminate theory. Then 3-point bending tests are conducted on the printed laminates to calculate their flexural stiffness. The influence of road/ fiber size and lamina layup on the flexural properties are also investigated. Furthermore, failure phenomena of printed laminates under bending loads is investigated.IntroductionIn recent times additive manufacturing (AM) techniques have risen in popularity [1] and these techniques have broad applications in the field of mechanical, aerospace and biomedical engineering [2].

AM techniques lack the limitations on fabricating parts with complex geometry in contrast to subtractive or conventional machining processes. However, the maximum size of the part remains a limiting factor when building using these advanced fabrication methods at present. The metal additive manufacturing industry has already found a way in the mass production of jet engines parts [4]. In near future, these AM processes will responsible for a major change in the design of materials as well as components. Traditionally, parts have been designed with manufacturability being a major consideration. Because of the limitation on complexity, the geometry of constructed parts was limited to what was reproducible using traditional manufacturing methods. Modern AM techniques allow for manufacturers to tailor the material properties by designing the microstructure of a material. The material properties of the final printed part differ from initial material properties.

The material properties of the final printed part differ from initial material properties. The change in the mechanical properties of a material of the printed parts is addressed in [5]. The difference in the material properties is due to changes in the microstructure of a material that is taking place during layer by layer printing while building a 3D part. The variation in the mechanical properties should be taken into consideration when measuring the final properties, which are useful in the design and analysis of the parts.Fused deposition modeling is an AM technique, in which a 3D part is fabricated by deposition of material one layer at a time. FDM is one of the most popular and widely used processes among all additive manufacturing techniques for 3D printing of parts. In the FDM process, hot molten filament material is extruded from the nozzle and deposited on a bed. The filament on the bed is referred to as the road.

In the FDM process, hot molten filament material is extruded from the nozzle and deposited on a bed. The filament on the bed is referred to as the road. Subsequent roads are deposited side by side and will have a bond with previously deposited road to form a single layer, then the next layer is deposited on the previous layer to build a 3D part. Each layer in a printed part behaves as an orthotropic material even though the initially filament material is isotropic [7]. The layers can be treated as a unidirectional fiber reinforced lamina in laminated composite structure. This means, the roads in a layer act as fibers in a lamina and several layers together with different printing orientations behave as laminated composite structures [16]. Therefore, parts printed via FDM can be considered laminated composites for the purposes of the characterization of their mechanical behavior. In this paper, the mechanical behavior of the printed laminated parts under bending load is investigated.

In this paper, the mechanical behavior of the printed laminated parts under bending load is investigated. Researchers [6-8] have been working towards the characterization of mechanical behavior of printed parts. It is revealed that the process parameters of the FDM process influence the mechanical properties by researchers [9-11] and further these are cause for introducing anisotropy in the printed part. The effect of different process parameters such as printing direction, layer thickness, gap between the adjacent fibers on the mechanical properties is investigated by [12]. It is found that the printing direction has a profound effect on the final properties of the printed part. The printing chamber temperature influences the bonding strength between adjacent fibers in992the part while printing [13-15]. Additionally, the boding strength between the roads has an effect on the final properties of the part.

Additionally, the boding strength between the roads has an effect on the final properties of the part. Kulkarni and Dutta [16] adopted the laminate theory for the characterization of FDM printed parts and then researchers [17,18] further worked on implementing the CLT for the FDM parts. The mesostructure of FDM printed parts influences the mechanical properties of the part. Alaimo et al. [19] carried- out experimental investigation to study the influence of the mesostructure and the composition of the filament on the mechanical properties on FDM printed parts. Furthermore, the CLT and laminate failure theory was adopted for studying the mechanical behavior of the printed parts. The filament material used in FDM process for printing parts is typically plastic; mainly PLA and ABS.

The filament material used in FDM process for printing parts is typically plastic; mainly PLA and ABS. Since the plastic material does not have better mechanical properties when compared to metals, researchers are working to improve the properties of FDM printed parts by adding reinforcements in the plastic filament material [20]. Composite filaments using carbon nanotubes as additional reinforcements in the thermoplastic materials have been developed [21,22] for better mechanical and electrical properties. These composite filament materials are broadening industrial applications of FDM processes, increasing adoption in for example biomedical industries [3].The FDM printed parts behave in an anisotropic nature, therefore the failure in such parts is complex. Failure phenomena of FDM printed parts under tensile, compressive and bending loads should be investigated. Ziemian et al.

Failure phenomena of FDM printed parts under tensile, compressive and bending loads should be investigated. Ziemian et al. [23] studied influence of printing direction/raster orientation on the mechanical properties and conducted failure analysis of FDM printed parts under tensile, compressive, flexural and impact loads. Furthermore, damage analysis of FDM printed parts under fatigue load is investigated in [24]. Hart and Wetzel [25] studied the fracture behavior of the fused deposited parts, where brittle behavior and ductile response were observed. They found that the facture behavior depends on the lamina orientation.The present paper mainly focuses on the flexural behavior of the laminates fabricated via fused deposited modeling. The elastic moduli in stiffness matrix of printed lamina are calculated by accounting the mesostructure of the FDM processed laminates in the finite element models.

The elastic moduli in stiffness matrix of printed lamina are calculated by accounting the mesostructure of the FDM processed laminates in the finite element models. For this, three different types of FE models are developed and simulated to find the elastic moduli. Then the elastic moduli of lamina are used in the classical laminate theory flexural behavioral characterization of the laminates. Also, flexural stiffness of the laminates is calculated for four different lamina layups. Then the 3-point bending tests are conducted on the printed laminate to find the flexural stiffness of the laminates. The stiffness values obtained from experiments are validated with analytical results obtained using CLT. Furthermore, the failure surface of cross ply laminates is characterized.

The stiffness values obtained from experiments are validated with analytical results obtained using CLT. Furthermore, the failure surface of cross ply laminates is characterized. Failure modes happened under bending loads that are discussed.MethodologyParts printed via FDM resemble unidirectional fiber reinforced laminated composite structures, therefore mechanics of laminated composites are adopted for the analysis of the printed laminates. The behavior of printed laminate with different printing direction in each layer is an anisotropic. Furthermore, the layer/lamina in the laminate behaves as an orthotropic material. The material behavior of the laminate is accounted for in the analysis by considering the constitutive matrix of the lamina in it. Therefore, the stiffness/constitutive matrix of lamina is required to find to the material behavior of the laminate in the characterization of their mechanical behavior.

Therefore, the stiffness/constitutive matrix of lamina is required to find to the material behavior of the laminate in the characterization of their mechanical behavior. This section initially describes the FE procedure to calculate the stiffness matrix of lamina. The stiffness matrix is calculated from FE simulations. Then the following sections after that explain analytical flexural behavior characterization of the printed laminates using classical laminate theory and finally experimental characterization is done.Constitutive matrix of a lamina: NumericalThe FDM process fabricates the parts by layer upon layer deposition of material, each layer acts a unidirectional fiber reinforced lamina. The lamina is a thin layer and therefore it can be treated as plane stress case.

The lamina is a thin layer and therefore it can be treated as plane stress case. The strain-stress relation for the lamina for a plane stress case is written as993ε  11  ε  22  γ  12    =    S 11 S 12 0S 12 S22 000 S66σ   11   σ   22   τ   12    Numerical suffix denotes coordinate system, 1 is along the fiber and 2 is across the fiber orientation.

That means the number of elastic constants of orthotropic material for a plane stress case reduced to four. These elastic constants are required to characterize the mechanical behavior of the FDM processed parts, which resemble laminates. These unknown constants can be obtained by conducting finite element simulation on three different tensile test specimens, explained below. All tensile test specimens considered in this paper follow standards mentioned earlier and elasto-static analysis is carried-out initially on these specimens using the finite element method.Let us consider a first case, where the fiber (road) is oriented along the axis of the loading in test specimen. The test specimen is subjected to uniaxial tensile loading, as shown in Fig. 1a. The elastic modulus E1 and ν12 of lamina can be calculated from this case.

The test specimen is subjected to uniaxial tensile loading, as shown in Fig. 1a. The elastic modulus E1 and ν12 of lamina can be calculated from this case. For a uniaxial tensile testing case, using stress-strain relation and strain- displacement relation, the strain energy is written asU=1 2E 1δ l  l   2VThe fiber orientation and type of loading in this case is replicated in the finite element model of the tensile test specimen after which the simulation is carried out. The strain energy (U) and displacement lδ along the axis of the load are obtained from finite element simulation. The length between grips l and volume of member that 1E is calculated from Eq. 5. The Poisson’s ratio given undergone deformation V are given. The only unknown, 22ε is lateral strain.

The length between grips l and volume of member that 1E is calculated from Eq. 5. The Poisson’s ratio given undergone deformation V are given. The only unknown, 22ε is lateral strain. The lateral strain is calculated by later displacement upon the original , where as 12 ν width in gage area of tensile member.= −ε ε 22 11Now a second case, where the fiber is oriented across the axis of loading in the test specimen and is subjected to uniaxial tensile load as shown in Fig. 1b. From this test the unknown elastic modulus E2 of the lamina is calculated. The total strain energy in this case is given as994,(1)(2)(3)(4),(5)U=1 2E2δ l  l   2VLet us take a case where the fiber is oriented off axis at 45° to the axis of loading in the tensile test specimen. The uniaxial load is applied to test specimen as shown in Fig. 1c. This test is conducted to find the unknown shear modulus G12 of the lamina.

The uniaxial load is applied to test specimen as shown in Fig. 1c. This test is conducted to find the unknown shear modulus G12 of the lamina. The total strain energy for this case is obtained from area under stress-strain curve and is given byU=1 E 2 xδ l  l   2VThe fiber orientation in the tensile specimen is replicated in the FE model and the strain energy (U) and xE is calculated from Eq. 7. displacement The relation between the principal 1-2 coordinates and non-principal x-y coordinates established through the transformation matrix, which is useful in this case to calculate the shear modulus G12 and the results is given bylδ along axis of the load are obtained from simulation.

The unknownG 12=4 Ex−1 E 11−1 E2+ν 2 12 E 1For further details about the procedure for calculation of elastic constants of the unidirectional fiber reinforced lamina refer to [27].Summary of linear finite element simulation procedure:a) Develop three CAD models of the tensile test specimen with three different mesostructures (3 casesmentioned earlier, shown in Fig.1) that would be obtained from FDM process.b) Mesh the CAD models with three dimensional finite elements. c) Apply boundary and load conditions to simulate the tensile testing of each case. d) Carry out the linear finite element analysis to find the strain energy and displacement that would beobtained from the simulation of each case. Gν , , 12 12e) Calculate the elastic moduli ( 1E E ,2) of the lamina from three cases using Eqs. 5, 6 and 8, asexplained above.Figure 1.

Gν , , 12 12e) Calculate the elastic moduli ( 1E E ,2) of the lamina from three cases using Eqs. 5, 6 and 8, asexplained above.Figure 1. Dimensions of tensile specimens (a) fiber orientation along x-axis, (b) fiber orientation along y-axis, (c) fiber orientation 45° to x-axis.995(6)(7)(8)(a)(b)Figure 2. Mesostructures (a) mesostructure of printed part via FDM (b) finite elements modeling of mesostructureThe strain energy is the energy stored by a member during the deformation and this can be calculated using finite element analysis. The calculation of strain energy of the FDM processed tensile test specimen subjected to uniaxial load is explained here and it is useful for finding the elastic constants of the layer(lamina). The uniqueness of this analysis is that the mesostructure of the material of the part that is obtained from FDM is replicated in finite element model of the tensile test specimen.

The uniqueness of this analysis is that the mesostructure of the material of the part that is obtained from FDM is replicated in finite element model of the tensile test specimen. This is how one can realistically simulate the FDM-processed materials to capture their mechanical behavior from the microscale level. The tensile test specimen and its cross- section with internal mesostructure obtained from 0° raster orientation to the x-axis in FDM are shown in Fig. 2a. The tensile test specimen is as per ASTM D3039 and it is subjected to uniaxial tensile loading, P along the x-axis. All dimensions of the tensile test specimen shown in Fig. 1 are in millimeters. Each road in a layer is treated as a fiber and is modeled using 3D linear finite elements as shown in Fig.2b. This means that the 3D constitutive relation for isotropic material is accounted in the finite element analysis.

This means that the 3D constitutive relation for isotropic material is accounted in the finite element analysis. Elasto-static uniaxial tensile tests were carried out in the commercially available finite element software, Hyperworks.(a)(b)Figure 3. Fiber orientation 0°/90° in a mesostructure of the printed part using FDM Stratasys µ print plus machine (a) 0.330 mm, (b) 0.254 mm layer thickness.The mesostructure of the printed part is shown in the Fig. 2a. The width of the road is larger than the height of the road, the height of the road is equal to the layer thickness (t). The road takes on an elliptical cross section during deposition and solidification of the material [17,26]. The dimension of cross section of the road is equal to the layer height and its major axis is double the layer height. Let us consider an example mesostructure with a layer height of 0.330mm. This mesostructure is replicated in the FE models of tensile specimens.

Let us consider an example mesostructure with a layer height of 0.330mm. This mesostructure is replicated in the FE models of tensile specimens. The specimens are subjected to uniaxial tensile load at one end and other end is fixed. In the present cases, the unit displacement (δl) as a tensile load along x axis is applied at one end. Then finite element simulations are carried out on the three tensile models. The strain energies for three cases are taken from the finite element simulations. The total strain energy of the tensile specimen in which the fibers are along the x-axis is 271.68 N-mm. 462.06 N-mm is the strain energy for the case where the fibers oriented along the y-axis. The strain energy of the specimen for the last case is 213.72 N-mm. The elemental strain energy contour plots for the three cases are shown in Fig.4. The elastic moduli are calculated by inserting strain energies in the equations 5, 6 and 8, and are presented in the Table 1.996(c) Figure 4.

The elastic moduli are calculated by inserting strain energies in the equations 5, 6 and 8, and are presented in the Table 1.996(c) Figure 4. Strain energy plots of the uniaxial tensile specimens (a) fibers oriented along the load (b) fibers oriented transverse to the load (c) fibers off axis to loadThe mesostructures of the printed parts using Stratasys µ print plus printer for two different cases are shown in Fig.3. Similarly, a mesostructure with a layer height of 0.254 mm is replicated in the FE models of the tensile specimens and then simulations are carried out to calculate the elastic moduli for this case. The elastic moduli for this case are presented in the Table.1. The present numerical results are compared with existing experimental results. The difference between the experimental and present results could be due to the variance in mesostructures present in the experimental tensile coupons by the researchers.

The difference between the experimental and present results could be due to the variance in mesostructures present in the experimental tensile coupons by the researchers. However, the present numerical results are comparable with results of the experiments. The elastic moduli for layer thickness 0.330 and 0.254 are nearly same. This means that the layer thickness on elastic moduli is not greatly influenced in the present analysis. However, the elastic moduli in direction 2 could be improved if the layer thickness is further reduced. The reason for this is because the volume of the void areas between the adjacent roads would be reduced.Table 1.

The reason for this is because the volume of the void areas between the adjacent roads would be reduced.Table 1. Lamina properties of the FDM processed part Elastic Moduli 11E , MPa 22E , MPa 12ν 12G , MPaExperimental Li et al [17] Rodriguez et al [9] Ziemain et al [24]2030.9 1251.6 0.34 410.01972 1762 0.37 6761820.5 1103.2 - -Present Numericalt =0.330 mm 2205.1 1456.3 0.34 673.9t= 0.254 mm 2205.7 1485.4 0.34 651.1The elastic moduli of a layer obtained from the previous section would be useful in the constitutive matrix of a lamina for the analysis of laminated composite parts. In the next section, the elastic moduli are used in the characterization of the flexural behavior of the 3D printed parts using classical laminate theory.997(a)(b)Classical laminate theory for flexural behavior study: AnalyticalIn FDM, the filament is melted in the extruder and the molten liquid is pushed through the nozzle to deposit on the substrate in the x-y plane.

Then the molten material on the substrate cools and solidifies into what is referred as ‘road’. Depositing a number of roads side by side forms a single layer. Several layers together act as a laminate. Such fabricated parts are considered as unidirectional fiber reinforced laminates. A two-layer plate fabricated with FDM process with an orientation of the roads in layer 1 and 2 at 0° and 90° to x-axis of the plate, respectively is shown in Fig. 3.The process parameters of the FDM influence the mechanical properties of the printed part. The raster orientation while depositing molten material in each layer can be defined by the user. Different orientations of road in layers lead to anisotropy in the properties of the printed parts. Process parameters such as the gap between two adjacent roads and the percentage of infill while printing governs the strength of the part.

Process parameters such as the gap between two adjacent roads and the percentage of infill while printing governs the strength of the part. The internal architecture of a part fabricated using FDM is not significantly different from that of fiber reinforced laminate structure. Therefore, the laminate theory for analysis of composite laminate can also be used for analysis of FDM parts. The FDM processed parts resemble laminate structures and therefore these parts can be characterized with classical laminate theory. The constitutive relation for a lamina is available in Eq.8 and is rewritten hereσ  11  σ  22  τ  12    =    Q 11 Q 12 0Q 12Q22 000Q66ε   11   ε   22   γ   12    The global coordinate system (x,y,z) for a laminate plate and local coordinate system (1,2,3) for a lamina are considered.

14 and 15 become{ } N=N∑k= 1 Q    h k+ 1∫h k{ } 0 εdz+h k+ 1∫h k{ } k z dz   (16){ } N=[A]{ } [ 0 ε +B]{ } k(17){M}=N∑k= 1 Q    h k+ 1∫h k{ } 0 εz dz+h k+ 1∫h k{ } 2 k z dz   (18){M}=[B]{ } [ 0 + εD]{ } k(19)where Nxx and Nyy represent the normal forces in the x and y directions (per unit width); Nxy is shear force; M xx and M yy denote the bending moments in the yz and xz planes; M xy is the twisting moment; [A], [B] and [D] are the extensional stiffness matrix, coupling stiffness matrix and bending stiffness matrix for the laminate, respectively. The matrices [A], [B], [D] are functions of each lamina stiffness matrix [Q] and the distance from mid plane of the laminate to the laminas.The mid-plane strains and curvatures can be calculated from Eq. 17 and 19, once we know the normal force and moment acting on a lamina.

17 and 19, once we know the normal force and moment acting on a lamina. A symmetric laminate layup will have identical plies (including material, thickness and orientation) located at an equal distance above and below midplane of the laminate. For a symmetric laminate coupling matrix, [B]=[0] and therefore there is no extension-bending coupling. Therefore, the strains for a symmetric laminate subjected to only transverse loads are given from Eq.17 as    kk kxxyyxy    =[D] 1−    MM Mxxyyxy    (20)In the 3-point bending test, the load is applied in the z direction and for laminate thickness h, M xy=0. The relationship between flexural stress and stiffness is written as the flexural modulus of elasticity of the laminate along the x direction is given as followsf E σ ε xxf xf xx=xxM ≠ , M yy=0 and 0 , using these above relationsf xE=[D12 − 1 ] 11h3(21)fxE of the laminate is calculated from Eq.

21, it is required to have the lamina’s elastic moduli For instance, if the such as E1, E2, G12, ν12. These constants are found from the FE simulation of tensile testing mentioned above. xE of the laminate calculated using Eq. 21. To Using these constants, the matrix [D] can be obtained, and then xE from the load versus validate the calculatedfffxE of the laminate, experimental work is carried out to find999deflection curve of the bending test. The flexural stiffness of four different lamina layups is calculated using Eq. 21 and are presented in Table 3.Flexural analysis of printed laminates: ExperimentalThis section describes about the flexural behavior of the printed laminates. The influence of fiber orientation and fiber thickness on the flexural stiffness is also studied using 3-point bending test. The FDM processed parts behave as a composite laminate structure, therefore mechanics of the composite laminates can be employed in their analysis.

The FDM processed parts behave as a composite laminate structure, therefore mechanics of the composite laminates can be employed in their analysis. The flexural strength and stiffness of the laminate depends on several factors such as fiber orientation and type of lamina layup. Therefore, the influence of such factors on the flexural stiffness of the laminates should be investigated. The delamination, one of the major failure mode, in the laminates is due to bending loads. Therefore, it is important to study the flexural characteristics of the FDM processed parts for reliable design. In the present experimental investigation, a 3-point bending test, ASTM D7264 standard for the composite laminates is considered. The dimensions of the bending test specimen are shown in the Fig. 5. All dimensions of the specimen are presented in millimeters.

The dimensions of the bending test specimen are shown in the Fig. 5. All dimensions of the specimen are presented in millimeters. To perform the test, the specimen is placed on two supports and then quasi-static transverse load is applied to the testing coupon at the center, as shown in the Fig. 5.Figure 5. Bending testing of printed laminate and its dimensionsThe specimens are printed using ABS-P430 supplied by the machine supplier and are built on a Stratasys µ print plus with default process parameters. Here, four different lamina layup are considered for two different cases. The layer size is different in both cases; 0.330 mm is the layer height in case a and 0.254 mm in case b. The dimensions of the printed laminates are available in Table 2. For each lamina layup, six samples are printed on a FDM machine that means 24 for each case. The orientation of fiber in four different laminates are provided in Table 3. Then, bending test are conducted on all specimens.

The orientation of fiber in four different laminates are provided in Table 3. Then, bending test are conducted on all specimens. The load versus deflection curves obtained from xE ) and its standard the tests are provided in the Fig.6 for case a and Fig.7 for the case b. The flexural stiffness ( deviation (SD) are calculated using standard equations available in the ASTM D7264 standard. The experimental results for the case a and case b are provided in the Table 4 and 5, respectively.fTable 2. Printed laminate dimensionsCase. a Case. bFiber size, mm Laminate thickness (t), mm0.330 0.2544.3 4.3Number of layers 13 17Span length, mm 137.6 137.6Table 3. Elastic constants of a layer and symmetric lamina layup of the tensile specimenElastic moduli of a layer for Case. a,Elastic moduli of a layer for Case.

Load-deflection curves for a thin layer laminates (a) Layup 1, (b) Layup 2, (c) Layup 3, (d) Layup 4.Thinner layer laminates have higher loading capacity when compared to thick layer laminates. Furthermore, these thinner layer laminates possess higher flexural stiffness than that of thick layer laminates. However, both thick and thin layer laminates have nearly same maximum deflection. This shows that the thickness of the fiber/road size or layer height influences the flexural stiffness, loading bearing and deflection of the laminate. The printing direction/fiber orientation of a lamina in the laminate also affected its flexural behavior. The flexural stiffness of the four different lamina layups calculated from CLT and experiments are provided for thick and thin layer laminates in the Tables 4 and 5, respectively. The flexural stiffness is higher for the case of layup 4 and lower for the layup 1 of laminate.

The flexural stiffness is higher for the case of layup 4 and lower for the layup 1 of laminate. Also, the layup 4 is undergone more deflection before failure when compare to the layup 1. For the thin layer laminates, some of the test coupon of the layup 4 laminates are not failed even at higher deflection. The flexural stiffness of the layup 4 type laminates is higher for the printed parts in contrast to the stiffness calculated using the classical laminate theory. The trend of stiffness of the laminate layup 1 to 3 calculated from experiments and CLT is decreasing, but for the laminate layup 4 is higher for the printed parts compared to the CLT results.Table 4. Flexural modulus of elasticity of the laminate from classical laminate theory for thick layers Ef x , MPa Layup 1 Layup 2 Layup 3 Layup 4Analytical, CLT 1923.1 1888.5 1814.7 1754.5Experimental ±SD 1942.3±32 1867.9±53 1779.4±40 1903.5±42Difference, % 0.98 -1.10 -1.98 7.821001Table 5.

Flexural modulus of elasticity of the laminate from classical laminate theory for thin layers Ef x , MPa Layup 1 Layup 2 Layup 3 Layup 4Analytical, CLT 1914.9 1870.6 1782.0 1723.2Experimental ±SD 2074.0±34 2013.5±43 1997.8±40 2069.8±29Difference, % 7.67 7.09 10.8 16.7Failure analysis of 3D printed coupons under bending load:The failure phenomena observed in the printed laminates is the fiber breaking and debonding of adjacent fibers. A laminate subjected to pure bending load and its crack surface are shown in the Fig.8. Failure mode in the laminate is ductile failure and the laminate layup 4 has higher ductility when compared to the laminate layup 1. The fiber breakage seen in a lamina, in which fibers are oriented at angle or along the x-axis and debonding is also seen in subsequent lamina, in which fibers are printed perpendicular to the orientation of fibers in previous lamina.

Therefore, the crack surface is perpendicular to the fibers in the top or bottom surface layer of a symmetric laminate. The failure surface of a cross ply laminate is shown in Fig.8c. The layers of a laminate subjected to pure bending load and undergo tensile stress, if they are below the mid-plane and compressive load if they are above the mid plane. The load distribution in the fibers of a cross ply laminate is shown in Fig 8b., the fibers in bottom layers subjected to tension (T) and the fibers on top surface subjected to compressive (C) load. In case of the laminate layup 4, the fiber breaking is seen in all layers of the laminate. This is because the fibers in each lamina are subjected to same quantity of load, where as in other lamina layups the fibers oriented in layers along or at an angle to the length of the test coupon are subjected to higher tension load and the fibers in the subsequent layer the same laminate are subject to lower tension.Figure 8.

Printed laminate subjected to three point bending load (a) 0°/90° laminate (b) load on the fibers in a laminate (c) failure surface of cross ply laminate.ConclusionThe parts fabricated via FDM processes are treated as laminated composite structures. The mechanics of laminated composites are adopted for the chracterization of mechanical behavior of the parts. The finite element procedure is presented to calculate the elastic moduli of a lamina of a printed laminated via FDM. The elastic moduli are used in the stiffness matrix for calculating the flexural stiffness of a printed laminate using classical laminate theory. Then the laminates of four different lamina layup with different layer thickness are printed. The effect of layer thickness and layup/printing oreintation on the flexural bhavior is investigated nummerically and experiemtnally.

The effect of layer thickness and layup/printing oreintation on the flexural bhavior is investigated nummerically and experiemtnally. Thin layer printed laminates have maximum load bearing capacity and also higher energy absorption when compare to that of thick layer laminates. The FE procedure presented in this study can be used to find stiffness matrix of a printed lamina. Furthermore, it is confirmed with the experimental results that the1002classical lamiante theory can be adopted for the characterization of mechanical behavior of the FDM printed laminates.References[1] Bourell, D.L., 2016. Perspectives on Additive Manufacturing. Annual Review of Materials Research, 46, pp.1- 18.[2] Guo, N. and Leu, M.C., 2013. Additive manufacturing: technology, applications and research needs. Frontiers of Mechanical Engineering, 8(3), pp.215-243.[3] Wang, X., Jiang, M., Zhou, Z., Gou, J. and Hui, D., 2017. 3D printing of polymer matrix composites: A review and prospective.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceA Floor Power Module for Cooperative 3D PrintingJacob Currence, Rolando Morales-Ortega, Jason Steck, Wenchao ZhouThe AM3 Lab, Department of Mechanical Engineering, University of Arkansas at Fayetteville Fayetteville, AR, United States of America; Corresponding email: zhouw@uark.eduAbstractCooperative 3D printing is an emerging technology that utilizes multiple printhead- carrying mobile robots to work simultaneously to 3D print or assemble products on a factory floor, which can provide the scalability, increased printing capability, and reduced human intervention for 3D printing to potentially become a mainstream digital manufacturing technology. However, powering the mobile printers for them to span entire factory floors poses an issue.

However, powering the mobile printers for them to span entire factory floors poses an issue. Traditional cords are not an option due to restricting free movement across long distances. On-board batteries would waste energy due to additional weight and the need to recharge could interrupt ongoing print jobs and increase printing time. In this paper, we present an electrified floor to power the mobile printers wirelessly. First, we designed a floor module with stainless steel conductive strips in a concrete base and a brush that is carried by the mobile robots to make sure it never loses contact with the electrified floor while in motion. Then we designed a circuit to sort the polarity of the current from the floor based on the power requirements of the robot. A prototype of the floor power module was then developed and tested with a mobile 3D printer. Results show the developed floor power supply can power the mobile 3D printers effectively.

A prototype of the floor power module was then developed and tested with a mobile 3D printer. Results show the developed floor power supply can power the mobile 3D printers effectively. This development will potentially enable an autonomous factory equipped with thousands of mobile 3D printers powered wireless by the factory floor.Keywords: Additive manufacturing; cooperative 3D printing; power system; mobile robot1. Introduction3D printing has quickly grown from a rapid prototyping tool to a serious contender of digital manufacturing over the past three decades [1-4]. Several major issues are still plaguing this technology for it to become a mainstream digital manufacturing method, including the limit in maximum print size, long printing time, printing complex products, and printing quality [5, 6].

Cooperative 3D printing is an emerging 3D printing technology that employs a plurality of mobile robots carrying different printheads to 3D print and assemble products on a factory floor. By freeing the printhead out of a “box”, the mobile 3D printers will be able to print over the entirety of factory floors and cooperate with each other to provide the scalability in both print size and print time while also enhancing the printing capability and quality by carrying different types of “printheads” (e.g., a gripper for incorporating pre-manufactured components, an inkjet printhead for high resolution printing, etc.). However, powering these mobile robots wirelessly poses an issue for this technology.The power source of a mobile robot is often the limiting factor for its maximum operating time without interruption, as well as its ultimate maneuverability and effective working space.

Within the manufacturing industry, there has been significant work done in developing power sources for mobile robots, ranging from rechargeable/replaceable batteries [7-10], hybrid fossil1661fuel-electric motors [11-13], and various solar, fuel cell, and hydraulic-electric power systems [14- 19]. Batteries are by far the most common form of mobile power storage, which can be a valid candidate for powering a mobile 3D printer. Battery-powered systems are highly mobile, and it is essential that a 3D printer is operated indoors, which makes it difficult to use solar or combustion- electric power sources. However, batteries are heavy to carry and must be accompanied by an additional power source or charging station if they are to operate for long periods of time on a single charge.

However, batteries are heavy to carry and must be accompanied by an additional power source or charging station if they are to operate for long periods of time on a single charge. Solutions include the onboard AC generator on the AURORA agricultural robot [20], charging stations utilized by KIVA Systems factory robots [21], and a tethered power cable such as in the HYPOS tethered power supply [8]. The fundamental requirement of a mobile 3D printing power source is that operation is not interrupted during the printing process, in particular for a print with numerous cooperative mobile 3D printers. In addition, any wired power source poses issues with cord tension disturbing the delicate printing process or blocking the paths of the mobile robots.

In addition, any wired power source poses issues with cord tension disturbing the delicate printing process or blocking the paths of the mobile robots. Therefore, all aforementioned solutions are eliminated for application to the mobile robots for cooperative 3D printing, and there exists a distinct need for a novel wireless power source for indoor mobile robots requiring long-term uninterrupted operation.In this paper, we present an electrified floor to wirelessly power mobile 3D printers, which allows for a constant wireless power source operating independent of any wires or energy storage devices. This design draws inspiration from the design of bumper car power systems where each bumper car draws power through a direct connection with a powered floor [22]. The bumper car power system occasionally loses power and relies significantly on the momentum of the bumper car to restore connection with the power floor, while the mobile 3D printer must never lose power during the printing.

Besides, the scale of the geometry and power requirement is very different from what is needed for a mobile 3D printer, which calls for a different design. In this paper, we present a power floor to meet the requirements of cooperative 3D printing. First, we design a power floor with electrified stainless steel conductive strips of alternating polarity installed within a concrete base. A brush apparatus is then designed to be carried by the robots across the conductive strips, transferring power from the floor to the robot while in motion. A rectifier circuit is then designed to sort the polarity of the current from the floor based on the power requirements of the robot, such that current flows in the correct direction regardless of the polarity of the active strips. A prototype of the floor power module is then developed and tested with a mobile 3D printer. Results show the developed floor power module can power the mobile 3D printers effectively.

A prototype of the floor power module is then developed and tested with a mobile 3D printer. Results show the developed floor power module can power the mobile 3D printers effectively. This development will potentially enable an autonomous factory equipped with thousands of mobile 3D printers powered wirelessly by the factory floor.This paper is organized as follows. In section 2, the engineering requirements and design of a floor power module, brush apparatus, and rectifier circuit compatible with an existing mobile 3D printer are presented. Section 3 presents the prototyping process of the designs. Then testing and results are discussed in section 4. Conclusions are given in section 52.

Section 3 presents the prototyping process of the designs. Then testing and results are discussed in section 4. Conclusions are given in section 52. DesignRecently, a mobile robot equipped with a Fused Deposition Modeling (FDM) printhead has been developed in the AM3 Lab at the University of Arkansas for cooperative 3D printing, as1662shown in Figure 1, which envisions an autonomous factory equipped with many of these mobile robots that work cooperatively to print and assembly products.Figure 1. Mobile 3D printer design.Our objective is to design a power module that provides wireless, uninterrupted power to multiple mobile 3D printers for use in cooperative 3D printing, such that no onboard power sources are required (e.g., batteries). A design inspired by the bumper car powering system has been pursued. The floor surface on which the robots move and print consists of parallel conductive metal strips that are each separated by an insulator.

The floor surface on which the robots move and print consists of parallel conductive metal strips that are each separated by an insulator. These conducting strips alternate positive and negative charge, one after another. A brush apparatus is then designed which transfers power from the floor to the mobile 3D printer by making contact with at least one positive and negative strip at any time while mounted to the robot. To then sort the incoming polarity, a rectifier circuit is needed to ensure the current flows in the desired direction no matter the polarity of each brush. In this paper, the three components of the floor power module (i.e., floor surface, brush apparatus, and rectifier circuit) are each designed in sequence, and a schematic of the resulting system is shown in Figure 2:Figure 2.

Cross-sectional view of the floor power module and each component, where the polarity of the strips in contact changes while the direction of current delivered to the robot remains the same.16632.1 Floor Surface Design2.1.1 Design ConstraintsA series of engineering requirements were developed for the floor surface that wouldsuccessfully provide power to a mobile 3D printer:1. Conductivity: The conductive strips must have a low resistance to current. This is important due to the potentially large areas of floor transmitting power, thus making the resistance of each strip a main factor of the power module’s efficiency. Therefore, it is set that the floor must be able transmit 12 volts and 10 amps of power over distances greater than 5 ft with a voltage drop less than 2 V for our prototype design.2. Insulating strips: To couple with the conductive strip material, an insulating material must be chosen so that it completely insulates the conductive strips from one another.

Insulating strips: To couple with the conductive strip material, an insulating material must be chosen so that it completely insulates the conductive strips from one another. These are necessary to having adjacent conductive strips of opposite polarity.3. Durability: The floor must be able to withstand the electric and mechanical stresses put upon it by the robots brushes and wheels, and it must be able to withstand years of wear with little to no maintenance. This applies to both conductive and insulating strips. In case of damage, it must be easy to repair or replace at low cost.4. Short circuit prevention: To prevent short circuiting between the floors strips and the robots brushes, the insulating layer in between the strips must be uniform and wider than the diameter of a single brush. That is, a brush above the insulating layer must not touch both conductive strips (which would cause a short circuit).5.

That is, a brush above the insulating layer must not touch both conductive strips (which would cause a short circuit).5. Manufacturability: The materials must be practical at both the small and large scales, inexpensive, and easy to manufacture. The total cost of the floor should be kept under $10/ft2.6. Safety: the electrified floor must operate within safety limits for humans to walk on. 7. Traction: The floor must have a sufficient high friction coefficient to provide enough traction for the wheels of the robot such that it doesn’t slip.2.1.2 Conductive Strip Material SelectionAfter researching commercially available materials that are both conductive and strong, it was decided to use brushed 304 stainless steel, which has a conductivity of 2.39% of international annealed copper standard (IACS) [23].

To carry the needed voltage and current, we need to determine the resistance of the conductive strip, which can be calculated according to Equation 1, where R is the resistance, ρ is the resistivity of the material, L is the length of the strip, and A is the cross-sectional area of the strip:𝑅𝑅 = 𝜌𝜌𝜌𝜌/𝐴𝐴These dimensions must be balanced with the cost of material, while remaining compatible with the overall geometry of the floor and brushes. It was initially assumed that the width of the strips must be 1.5 in. Given the resistivity of 304 stainless steel is 7.2E-7 ohm-m [23] and the desired strip length of 5 ft, the thickness of each strip was calculated to be 0.018 inches thick when rounded to fit within the range of pre-manufactured 304 stainless steel sheets. In the future, these floors may be the size of entire factories, spanning hundreds of feet. The strips must be able to1664(1)carry the power long distances without losing much of its initial power.

The strips must be able to1664(1)carry the power long distances without losing much of its initial power. Given the dimensions of the strips and the power requirements, the voltage drop across each strip may be calculated using Ohm’s Law in Equation 2:𝑉𝑉𝑑𝑑𝑑𝑑𝑑𝑑𝑑𝑑 = 𝐼𝐼𝑅𝑅In regard to surface durability, 304 Stainless steel is very strong and corrosion resistant. With a hardness of B82 on the Rockwell B scale [24], it will not be easily worn by the sliding brushes. Moreover, its corrosion resistant properties will keep it free from rust and allow for easy cleaning. This alleviates concerns of the sliding brushes damaging the floor surface during operation.2.1.3 Insulating Strip Material SelectionA marine grade epoxy was then used for the insulating strips, as well as for the binding material between the steel strips and the concrete base. This material was chosen for its ease of manufacturing and replacement, along with sufficient durability and insulating properties.

This material was chosen for its ease of manufacturing and replacement, along with sufficient durability and insulating properties. The epoxy allows for a liquid to be poured into the void space between the steel strips, and after hardening they can be sanded flat, creating a seamless transition between strips. This installation process is described further in section 3 of this paper. Similar to stainless steel, marine epoxy is also corrosion resistant. Although the epoxy is softer than both the stainless-steel strips and the chosen brush material, it will still have a relatively long effective life due to low pressure and friction of the brushes upon it, and it is the cheapest and easiest to repair or replace out of all materials used in the floor power module.2.1.4 Cost AnalysisThe floor materials chosen are both relatively cheap and easy to obtain. 304 brushed stainless steel sheets can be purchased in bulk for approximately $2 per cubic inch.

304 brushed stainless steel sheets can be purchased in bulk for approximately $2 per cubic inch. This price drops significantly as bulk of purchase goes up. With the chosen thickness of 0.018 in, the stainless steel strips will cost about $5 per square foot of floor for small scale builds. The marine epoxy comes in at approximately $0.34 per cubic inch. For the design application of 0.018 inches plus the sanded excess, the cost of the epoxy is approximately $0.05/ft2. In the case of a 2500 square foot factory, the cost of materials to install our design of powered floor would cost less than $7500 as shown in Table 1.Table 1. Approximate relative costs of different floor sizes.Small to medium scale floor (4 ft2) Large scale floor (2500 ft2)Stainless Steel Cost Epoxy Cost $5/ft2$0.05/ft2$2.85/ft2$0.05/ft2Full Floor Cost $20.20$7250With the chosen thickness of 0.018 in, the stainless-steel sheets can be cut easily using common machinery.

On a larger scale, the strips may need to be pre-cut by the manufacturer. The1665(2)epoxy can be sanded down using industrial sanders on a small and large scale at relatively low cost. Therefore, it is possible to keep the total cost of the floor under $10/ft2.2.2 Brush Apparatus2.2.1 Design ConstraintsThe brush apparatus is carried by the mobile robot to draw power from the electrified floor. In a similar fashion to the design of the floor surface, a series of engineering requirements were taken into consideration during the design:1. Geometry: The brushes must be arranged in a geometric pattern that ensures at least one brush is in connection with a positive strip and another in connection with a negative strip. This must be the case at all times no matter the location or orientation of the robot on the floor. In addition, a single brush must not touch both a positive and negative strip simultaneously, as this would cause a short circuit.2.

In addition, a single brush must not touch both a positive and negative strip simultaneously, as this would cause a short circuit.2. Reliability: The brushes must keep constant contact with the floor in order to ensure constant power, including when the brushes encounter defects or debris. In particular, the brushes must not have issues crossing between each adjacent strip. For example, the brush must not get caught on the insulating layer as it slides across the surface.3. Dimensions: The brushes must be mounted underneath the robot in order to reach the floor. This leaves open only one spot to mount the brushes with the current mobile 3D printer design: a small central area under the robot between the stepper motors as can be seen in Figure 1. This equates to a rectangular prism 6x3.5 inch wide and 2.1 inch tall.4.

This equates to a rectangular prism 6x3.5 inch wide and 2.1 inch tall.4. Brush material: The material used for the brushes must be conductive, mechanically resilient (strength and hardness), and economically viable (low cost and ease of manufacturability). To prevent excessive scratching of the floor surface, the brush material must have a Rockwell hardness value within 5 of the chosen floor material.2.2.2 Brush GeometryThe mobile 3D printer will carry a set of brushes along the floor power module that allows the robot to switch between conductive strips without disconnecting from power. In order to provide continuous power, there must be at least one brush on each polarity of strip at all times. The more brushes in contact with the strips on the floor the more chances there are that one of them will lay on a positive strip and other one on a negative strip.

The more brushes in contact with the strips on the floor the more chances there are that one of them will lay on a positive strip and other one on a negative strip. However, since having an excess of brushes can be both expensive and impractical, it was imperative to reduce the number of brushes by strategically selecting their positions. The positioning of the brushes relative to each other is defined as the geometric configuration. A feasible geometric configuration for the brushes will be determined below.2.2.2.1 Two BrushesGeometric analysis begins by establishing the constraint that at least two brushes are required at all times, one for positive and one for negative. If there are only two brushes on the power module, these are free to move vertically from end to end while providing continuous1666power. However, horizontal displacement is restricted, since as soon as one brushes encounters an insulating strip, the circuit is disconnected.

However, horizontal displacement is restricted, since as soon as one brushes encounters an insulating strip, the circuit is disconnected. This event is illustrated in Figure 3. This indicates that two brushes are not sufficient for powering the circuit.Figure 3. Illustration of two brushes at the moment of disconnection during horizontal movement. The dotted circles represent the brush locations before movement and the solid circles represent the brush locations after the movement.2.2.2.2 Three BrushesBy then adding a third brush to the configuration, the issue with horizontal movement can be solved, as illustrated in Figure 4(a). As the brush apparatus moves across the floor surface, each brush transitions across an insulated strip without losing overall power. However, an equivalent issue arises when rotational motion is considered.

However, an equivalent issue arises when rotational motion is considered. It is found that as the robot transitions from horizontal to vertical, the horizontal distance covered by the brushes decreases until it is unable to cover the distance to touch at least one positive and negative strip. This issue is depicted in Figure 4(b), and suggests that the rotation of the brush geometry must place one brush on either of the farthest sides of the geometry, such that both strips are in contact.1667Figure 4. Geometry diagram with three brushes: (a) translational motion; (b) rotational motion.2.2.2.3 Four BrushesFrom the previous arrangement, it is clear that a geometric configuration with at least four brushes is required to achieve continuous power supply when the robot is subjected to both translational and rotational displacements.

By positioning three brushes in an equilateral triangle, and by placing the fourth brush at the center of a circle intersecting each vertex of the triangle, a geometric configuration for four brushes is obtained. This configuration has four critical positions, as shown in Figure 5(a). To analyze potential disconnections during translational displacements, the position of each brush is projected onto the horizontal in Figure 5(a). These horizontal positions are then translated across the floor surface in Figure 5(b). In the same fashion as Figure 4(a), each brush transitions through positive, disconnected, and negative while there is at least one brush on positive and one brush on negative, thus maintaining a complete circuit throughout the movement. However, it is still necessary to identify disconnections during rotation. By observation, no degree of angular displacement causes loss of contact with positive and negative simultaneously.

However, it is still necessary to identify disconnections during rotation. By observation, no degree of angular displacement causes loss of contact with positive and negative simultaneously. For example, rotating clockwise the geometric configuration across the floor between positions 1 and 2, the rightmost brush loses contact with the positive strip while the leftmost brush is just beginning its positive motion (i.e., as one brush disconnects, the opposing brush is moving onto a strip instead of away, thus maintaining a closed circuit). This extends to every other position, since as one brush in transitioning, another is soundly connected the corresponding strip. Therefore, an equilateral triangle and central brush arrangement is determined to be a sufficient geometric configuration, and is then used for designing the brush apparatus.1668Figure 5.

Therefore, an equilateral triangle and central brush arrangement is determined to be a sufficient geometric configuration, and is then used for designing the brush apparatus.1668Figure 5. Four critical brush positions for equilateral triangle geometry: (a) Circular positions projected onto the horizontal plane; (b) translational motion of each configuration as each brush position transitions through positive, disconnected, and negative.2.2.2.4 Dimensional ProportionalityAfter determining that the equilateral triangle is a suitable geometrical configuration, it is necessary to determine the relations between triangle size and strip width.

The dimensions of the geometric configuration formed by the brushes and the width of the floor strips have a direct correlation (defined proportionalities) that make them interact with each other in a way that provides a constant transmission of electrical power from the floor to the main circuit.First, the size of the area covered by the robot will determine the size of the geometrical configuration since the brushes will be within the perimeter covered by the robot. The size of the triangle will always be smaller in size than the area covered by the robot to protect the brushes from other objects on the power module floor and avoid loss of electrical power. The dimensions of the floor surface and brush geometry are defined in Figure 6:1669Figure 6. (a) Dimensions of the floor surface and brushes; (b) brush geometry on floor surface in position 2 of Figure 5(a).The relations between each dimension can be determined by analyzing the moment the triangle is transitioning strips.

One critical insight is that the horizontal projection of the brushes must be present on 3 strips when either of the brushes at the corners is about the shift strips. Since the horizontal span of the brushes changes as the brush configuration rotates and the shortest horizontal span occurs when the brush configuration is rotated to position 2 (or position 4) as shown in Figure 5(a), we need to make sure the horizontal span is still sufficient to cover 3 strips. Therefore, a geometric relationship must be satisfied as shown in Equation 3:2𝑎𝑎 ≤ 𝐻𝐻 − 𝐷𝐷 − Equation 3 can be used to determine the maximum strip width for maintaining contact with 3 strips during a brush transition; however, issues arise if the strip width is too narrow, such as unintended brush transitions and skipping strips between brushes. Therefore, a minimum strip width is defined which, along with the maximum, establishes a working range for the conductive strip width.

Therefore, a minimum strip width is defined which, along with the maximum, establishes a working range for the conductive strip width. The minimum conductive strip width is determined by considering the distance between the central brush and any of the 3 remaining brushes in the geometric configuration, which is the vertex to center distance in the equilateral triangle. The summation of this distance with the brush diameter defines the minimum strip width. This constraint is illustrated by considering the horizontal projection of position 2 (Figure 5(b)) as brush 1 transitions strips, since if the strip width is too narrow, brush 2 will begin transitioning strips prematurely. This relation is quantified in Equation 4, where𝑏𝑏3 is the vertex to center distance:𝜌𝜌/√3 𝐿𝐿+ 𝐷𝐷 The final constraint is that the thickness of the brush must be thinner than the width of the insulating strips.

If the diameter of the brush is equal or larger than the width of the insulating strip,𝑎𝑎 ≥√1670(3)(4)a short circuit occurs. It is assumed that all brushes are of equal diameter, and this constraint results in the inequality in Equation 5:2.2.2.5 Determining Geometry Size𝑏𝑏 > 𝐷𝐷The space available under the robot has an area of 6 inch wide by 3.488 inch long. This means that the size to be selected for the geometrical configuration is constrained to this area. To keep the brushes within this area, H (see Figure 6(b)) was chosen to be 2.00 inch. Then from this dimension, the length of the sides of the triangle, L, were calculated to be:2𝜌𝜌 = 𝐻𝐻√3𝜌𝜌 = (2.00)2 √3𝜌𝜌 = 2.32 𝑖𝑖𝑖𝑖Increments of 1/32 inch were used to find the brush diameter and the width of the insulating strips.

It was decided that since the width of the insulating strips should be larger than the diameter of the brush, the width of the insulating strip should be the next increment of 1/32 inch. These values were introduced in Equations 3 and 4 simultaneously and plotted in Figure 7:Figure 7. Plot of the conductive strip width as a function of the brush diameter, along with the insulation strip width which corresponds to each brush diameter.1671(5)(6)From Figure 7, any combination where amax > amin meets the geometric requirements of the brush design. However, a combination of 1/8 inch and 5/32 inch for the brush diameter and the width of the insulating strips, respectively, were selected for this design, since increments of 1/32 inch are commercially produced.

Based on the calculation of the height H, the maximum and minimum size for the conductive strips can then be calculated by Equations 3 and 4, in addition to the brush diameter and insulating strip width determined via Figure 7:𝑎𝑎𝑚𝑚𝑚𝑚𝑚𝑚 ≤ 1.5725𝑎𝑎𝑚𝑚𝑚𝑚𝑚𝑚 ≥ 1.46Therefore, the width of the conductive strips should be between 1.5725 inch and 1.46inch, and the intermediate value of 1.5 inch was chosen as the final conductivestrip width.2.2.3 Brush Material SelectionBased on our design requirements, the brush needs to be able to carry at least 10 amps of current without causing a drop on the voltage. Assuming the maximum voltage drop allowed on the brush is 0.025 volt (0.2% of 12 volts), the maximum resistance of the brush can be calculated to be 0.0025 Ω. Based on the geometry described in the previous section, the diameter of the brushes was found to be 1/8 inch.

Based on the geometry described in the previous section, the diameter of the brushes was found to be 1/8 inch. Assuming the length of the brush is 1 inch, we can determine the resistivity of the brush material based on Equation 1:ρ=30 µΩ-inTherefore, a copper-tungsten alloy with a composition of 30% copper and 70% tungsten (W70Cu30) was chosen for the brushes, which has a resistivity of 28.34 μΩ-in [25]. In addition, the brushes need to be hard enough to withstand the constant friction forces, without scratching or damaging the floor. It is known that the copper tungsten alloy has a Rockwell Hardness (B) of B85-92 [26] while 304 stainless steel has a hardness of B82 [24]. These properties being similar in magnitude will make it very difficult for the brushes to damage the stainless steel, while still allowing the brushes to be easily machined to the proper length and shape.

Other candidate brush materials, such as copper and graphite, were shown to be either be too soft or brittle, or require a much larger brush diameter for this specific application.2.2.4 Brush Holder DesignThrough implementing the established geometry and dimensions, the brush apparatus was designed as shown in Figure 8. The final dimensions of this design are then presented in Figure 9.1672(7)(8)(9)Figure 8. (a) Labeled exploded view of brush holder assembly; (b) brush holder mounting brackets and wire inputs.Figure 9. Dimensions of the designed brush apparatus: (a) front view; (b) top view (all units in inches).To ensure constant connection with the floor despite any debris or defects, springs were mounted between the brushes and the robot. These springs apply a downward force on the brushes into the floor, ensuring a constant connection as the robot moves along the surface.

These springs apply a downward force on the brushes into the floor, ensuring a constant connection as the robot moves along the surface. Furthermore, if there is an inconsistency with the surface of the floor, the spring will allow the brush to rise or fall accordingly, and as the brushes wear down over time, more material is extruded as needed. This allows for less frequent maintenance of the robots. The springs must have enough strength to ensure a good connection, but not too strong as to cause unnecessary friction on the brushes.1673In order to ease assembly and increase strength, it was decided to mount all four brushes in a single housing that contains all of the components. This housing is rigid and consistent with the previously determined brush geometry. The housing is mounted directly to the underside of the robot using three simple screw and bolt attachment points.

This housing is rigid and consistent with the previously determined brush geometry. The housing is mounted directly to the underside of the robot using three simple screw and bolt attachment points. The top of the brush holder is designed to fit snugly around the X shaped supports underneath the bot (see Figure 1). This housing descends downwards towards the floor with four vertical hollow tubes that contain the springs and brushes, and allows the components to move freely inside. The housing is then capped off sealing the components inside the tubes, while allowing the brushes to extrude out of the housing and make contact with the floor. This cap is fastened using screws and bolts. The entire housing is suspended one millimeter off the floor, allowing the springs to push out the brushes and fill the difference.

This cap is fastened using screws and bolts. The entire housing is suspended one millimeter off the floor, allowing the springs to push out the brushes and fill the difference. A small gap was chosen to reduce the effect of linear and bending forces on the brushes by providing more support.2.2.4 Cost AnalysisFor future manufacturing on a larger scale, the brush material needs to be relatively cheap and easy to obtain. Copper tungsten alloy can be bought easily online and in bulk. The average cost of W70Cu30 tungsten copper alloy is around $17 per square inch, or roughly $8.5 per robot. These prices drop drastically when large quantities are bought in bulk. This is a good price for the application and quality of the material.2.3 Rectifier Circuit2.3.1 Design ConstraintsAs the mobile 3D printer moves across the floor surface, the alternating polarity received through the brushes must be sorted such that current flows in the correct direction for use onboard the printer.

A rectifier circuit was developed to accomplish this goal, and was subject to the following design constraints:1. Dynamic brushes: To keep cost and size to a minimum, a single brush must be able to switch from positive and negative current depending on its location. That is, the onboard rectifier circuit must separate and organize the charge no matter which brush it is coming from. For example, as long as one brush is positive and one brush is negative, the circuit forces the current to flow in the correct direction through the robot. This allows for a single brush to accept both positive and negative currents without causing a short circuit, thereby reducing the total number of brushes required.2. Power: The circuit needs to be able to handle the supplied voltage of 12 volts and 10 amps safely and without overheating.3. Four brush sorting: The circuit must sort the current of all four brushes simultaneously. 4.

Four brush sorting: The circuit must sort the current of all four brushes simultaneously. 4. Volumetric constraints: The entire circuit must be small enough to fit onto the robot without interfering with the existing electronics and mechanical systems.2.3.2 Circuit Design1674The rectifier circuit design shown in Figure 10 was created according to the specified engineering requirements. Each brush is connected to two Schottky diodes, in forward and reversed bias, making a total of 8 Schottky diodes in the rectifier circuit. Every time a brush is connected to a conductive strip, one of the diodes will be closed and the other will be open. This mechanism allows the current to always flow in the same direction through the rectifier circuit. In addition, there is an electrolytic capacitor connected in parallel to one of the ends of each of the diodes.

In addition, there is an electrolytic capacitor connected in parallel to one of the ends of each of the diodes. The electrolytic capacitor has the functionality of storing small amounts of charges during operation, thereby avoiding electric sparks when passing from one strip to the other. This minimizes the risk of fires or electrocution.Figure 10. Circuit diagram of the complete system, where each component is labeled by region.3. Prototyping3.1 Floor Surface3.1.1 Prototyping ConsiderationsA cement base was chosen because it would be non-conductive, is readily available, and would not deform under stress. In the future, the base could be any rigid non-conducting surface or floor. The cement was poured into a 2 x 2 ft. mold until approximately 1 inch thick, with carbon fiber rebar inserted for additional strength and crack prevention.

The cement was poured into a 2 x 2 ft. mold until approximately 1 inch thick, with carbon fiber rebar inserted for additional strength and crack prevention. Although the floor would not likely be experiencing wide temperature changes or tensile stresses, rebar was necessary so that no catastrophic damage occurred during transportation. Using cement demonstrated the viability of implementing this design on a large-scale factory floor.16753.1.2 Fabrication ProcessA 14 V power supply was used at this stage to prevent the possibility of a voltage drop across the prototype floor surface interfering with the performance of the mobile 3D printer, which requires a 12 V operation voltage. The floor surface was then fabricated using cement, 304 stainless steel strips, and marine grade epoxy through the following steps (corresponding to the steps depicted in Figure 11):Step 1: Cement was poured into a mold with a flat bottom until it is halfway full.

This becomes the base of the floor surface. Carbon fiber rebar was placed on top of the cement. Step 2: The rest of the mold was filled with concrete. Once dried the concrete base was removedfrom the wooden mold.Step 3: It is necessary that the spacing between strips is uniform across the floor. Two plastic spacers were designed and printed to fit over the ends of the strips while they were on the cement base. The spacers would maintain the gap between adjacent strips while the strips were adhered together using duct tape.Step 4: The strips were then removed from the cement base so that a very thin layer of marine epoxy can be spread over the surface. It is necessary to minimize the amount of epoxy used, as excess would result in humps and dips on the final floor surface. With the epoxy applied, the strips were placed back on the surface, and several weights were put on top to ensure even contact between the strips and the epoxy.

With the epoxy applied, the strips were placed back on the surface, and several weights were put on top to ensure even contact between the strips and the epoxy. The epoxy was left to cure for 24 hours.Step 5: Once the strips were fully adhered to the cement, epoxy was applied to the gaps between the metal strips. These would serve as the insulating strips on the floor surface. After the epoxy had cured, it was sanded down so the surface was flat. At this point, areas where the epoxy had not filled in the gaps, or where bubbles had formed were identified, and additional epoxy was applied and allowed to cure.Step 6: Each strip was then connected to a 14 V power supply, with each strip alternating positive and negative. Finally, the finished floor was housed in a recessed box to prevent inadvertent contact by users and to allow discrete mounting of the power supply. The final floor assembly was sanded down a final time, which yielded a flat surface.1676Figure 11.

The final floor assembly was sanded down a final time, which yielded a flat surface.1676Figure 11. Floor prototype construction process.3.2 Brush Apparatus3.2.1 Fabrication ProcessWith the fundamental design established, prototyping began with 3D printing the brush housing with an uPrint SE Plus 3D printer. 3D printing was the best option for prototyping, since it allows for prototypes to be made quickly and cheaply, and thus makes it possible to have many iterations early in the design process. The fabrication process followed was:Step 1: The copper-tungsten brushes (Figure 8(a)) were machined from a single 1/8 in diameter rod. First, the rod was cut by hand into 4 equal brushes of 0.5 in long.

First, the rod was cut by hand into 4 equal brushes of 0.5 in long. Then, the end that makes contact with the floor was machined to be smooth, and a small bevel was added to the edge to remove any sharp edges prone to snagging.Step 2: To wire the brushes, a small hole was drilled into the face of each brush facing the housing (Figure 8(a)), and wires were tapped into these holes and secured with superglue (cyanoacrylate).Step 3: The brush caps were then installed around each brush, which increases stability andprovides a platform on which each spring makes a solid connection.Step 4: The springs and brushes were loaded into the housing (Figure 8(a)). The wire was fed through the hollow tubes, through the center of the spring, and out through the wire outputs labeled in Figure 8(b).1677Step 5: The entire assembly was locked together by first attaching the housing cap to the main housing with screws.

Then, the complete brush apparatus was mounted to the base of the robot, as seen in Figure 12.Figure 12. Prototype of brush apparatus.3.3 Rectifier Circuit3.3.1 Component SelectionThe rectifier circuit is composed of a total of 8 Schottky diodes mounted on a printed circuit board (PCB) and arranged as is shown in Figure 13. The Schottky diodes were selected because of their great capability to withstand high frequency of change of polarities. A 1000 μF electrolytic capacitor was connected in parallel to each brush, as is shown in Figure 10. The PCB selected for the rectifier circuit is a 2 layer PCB with an area of 63.6 mm by 51.3 mm and a thickness of 1.02 mm. This double layer PCB has one layer of copper and other one of an insulating material.3.3.2 Fabrication ProcessThe layout of the circuit board was CNC machined and is shown in Figure 13.

This double layer PCB has one layer of copper and other one of an insulating material.3.3.2 Fabrication ProcessThe layout of the circuit board was CNC machined and is shown in Figure 13. The stepsof assembling the rectifier circuit are as follows:Step 1: The eight Schottky diodes, the 1000 μF electrolytic capacitor, the terminals of the four brushes, and the terminals of the positive and negative poles that feed to the 3D printer circuit are each mounted on the PCB in the orientation shown in Figure 9.Step 2: All electronic components were soldered by hand. Because of this, the spacing between the components was left wide enough to easily solder the components, and to allow sufficient room for the components to cool off during operation. Further studies will be carried out to determine the optimal spacing for the components to cool off or if additional heat sinks will be needed, and later models will be soldered automatically in a circuit printing machine.1678Figure 13.

Rectifier circuit prototype.4. Results4.1 Preliminary TestingThe floor power module was first tested by measuring if the conductive strips were consistently providing the necessary voltage. With the power supply connected to each strip alternating positive and negative, ideal performance would entail 14 V being read at any point across 2 opposite polarity strips. After the power supply was powered on, a voltmeter was used to test the voltage between strips. This measured a voltage of ~14.1 V at any point along the floor surface, verifying that the prototype was functional. This testing process may be seen in Figure 14:Figure 14. Strip voltage test setup.1679Next, a test was performed to show that the brushes could adequately and safely transfer the power from the floor to the mobile robot. A mock robot was constructed to test the floor power module prior to testing the mobile 3D printer to protect the real robot from any potential dangers.

A mock robot was constructed to test the floor power module prior to testing the mobile 3D printer to protect the real robot from any potential dangers. This mock robot is a replica of the real robot minus any hardware not necessary for testing the function of the brushes and circuit. The mock robot is shown in Figure 15. The mock robot is built upon an identical base to that of the mobile 3D printer, and contains all components of the floor power module. In addition, it contains a dummy load which simulates the resistance of the 3D printer in operation. This allows all components to be mounted in their intended arrangement.Figure 15. Mock robot on floor surface.For mock robot testing, the prototype for the brush apparatus was mounted to an identical base for the robot that will be used. Then, the brushes were connected to an Arduino equipped with a voltage and amperage meter in order to record and display the results.

Then, the brushes were connected to an Arduino equipped with a voltage and amperage meter in order to record and display the results. The floor was then fully powered and the mock robot with brushes equipped was moved around the floor by hand, accounting for every possible angle and position. Also, the mock robot was quickly lifted off of the floor to demonstrate what would happen in an unwanted case of disconnection. The results and corresponding path of movement can be seen in Figure 16:1680Figure 16. Mock robot test of floor power module: (a) current and voltage against time; (b) path of movement for mock robot during test.As desired, the robot received a constant 12 V and 10 A without any dangerous spikes or drops. When the robot becomes disconnected the power immediately drops to 0 W, and when connected it immediately regains full power.

When the robot becomes disconnected the power immediately drops to 0 W, and when connected it immediately regains full power. These are the results that were desired and designed for.4.2 Testing with Mobile 3D PrinterOnce confident in the safety of the floor and brushes, testing was continued with the real robot. The brush apparatus was mounted to the robot and wired to its onboard microcontroller. The robot was then placed on the floor surface and power was switched on, the robot immediately powered up. A G-code script was uploaded to the robot to perform a simple rectangular path along the floor to test connection and power. The robot ran without any interruptions, and this test can be seen in Figure 17. Other paths have been uploaded to the robot and they have been performed successfully as well. The results show the promise for a full implementation of this design in large scale in a factory setting.1681Figure 17.

The results show the promise for a full implementation of this design in large scale in a factory setting.1681Figure 17. Robot moving across the floor power module, without any additional power sources. Each image follows chronologically from the previous.5. ConclusionsThis paper presented a scalable floor power module to wirelessly power mobile 3D printers for a cooperative 3D printing platform. First, design requirements and constraints were established for each component. These requirements and restrictions guided the overall design to emulate bumper car powering technology, and it was established that a powered floor, brush apparatus onboard the printer, and a rectifier circuit must each be designed to achieve the goal of wirelessly powering mobile 3D printers. These components were then prototyped and tested by measuring the potential difference across the floor surface, followed by running a mock robot equipped with each component.

These components were then prototyped and tested by measuring the potential difference across the floor surface, followed by running a mock robot equipped with each component. Testing results showed that the mobile 3D printers can successfully be powered wirelessly by a conducting floor power module and a brush apparatus mounted beneath the robot. The intended impact of this design is to expand the overall printing envelope of 3D printing and to enable cooperative printing through the use of mobile printers by providing constant untethered power. By enabling cooperative 3D printing on a large scale, the technology of 3D printing advances towards being a competitive digital manufacturing method.6. AcknowledgementsWe gratefully acknowledge the financial support from the University of Arkansas, through the startup fund provided by the Vice Provost Office for Research and Economic Development.

AcknowledgementsWe gratefully acknowledge the financial support from the University of Arkansas, through the startup fund provided by the Vice Provost Office for Research and Economic Development. Any opinions, findings, and conclusions or recommendations expressed in this publication are those of the authors and do not necessarily reflect the views of the University of Arkansas.7. References[1] Billiet, T., Vandenhaute, M., Schelfhout, J., 2012, "A Review of Trends and Limitations in Hydrogel-Rapid Prototyping for Tissue Engineering," Biomaterials, 33(26) pp. 6020-6041. [2] Campbell, I., Bourell, D., and Gibson, I., 2012, "Additive Manufacturing: Rapid Prototyping Comes of Age," Rapid Prototyping Journal, 18(4) pp. 255-258. [3] Macdonald, E., Salas, R., Espalin, D., 2014, "3D Printing for the Rapid Prototyping of Structural Electronics," IEEE Access, 2pp. 234-242.

255-258. [3] Macdonald, E., Salas, R., Espalin, D., 2014, "3D Printing for the Rapid Prototyping of Structural Electronics," IEEE Access, 2pp. 234-242. [4] Weller, C., Kleer, R., and Piller, F., 2015, "Economic Implications of 3D Printing: Market Structure Models in Light of Additive Manufacturing Revisited," International Journal of Production Economics, 164pp. 43-56. [5] Oropallo, W., and Piegl, L., 2016, "Ten Challenges in 3D Printing," Engineering with Computers, 32(1) pp. 135-148. [6] Gao, W., Zhang, Y., Ramanujan, D., 2015, "The Status, Challenges, and Future of Additive Manufacturing in Engineering," Computer-Aided Design, 69pp. 65-89. [7] R. C. Luo, C. H. Huang, and C. Y. Huang, 2010, "Search and track power charge docking station based on sound source for autonomous mobile robot applications," 2010 IEEE/RSJ International Conference on Intelligent Robots and Systems, Anonymous pp. 1347-1352.

1347-1352. [8] Wattanasin, C., Aiyama, Y., Kurabayashi, D., 2001, "Hybrid Power Supply for Mobile Robots," Advanced Robotics, 15(6) pp. 695-710.1682[9] Wu, J., Qiao, G., Ge, J., 2012, "Automatic Battery Swap System for Home Robots," International Journal of Advanced Robotic Systems, 9(6) pp. 255. [10] Youn-Ho Choi, and Dong-Ha Lee, 2013, "Battery swap station for mobile robot," 2013 International Conference on Electrical Machines and Systems (ICEMS), Anonymous pp. 1445- 1447. [11] J. B. Siegel, Y. Wang, A. G. Stefanopoulou, 2015, "Comparison of SOFC and PEM Fuel Cell Hybrid Power Management Strategies for Mobile Robots," 2015 IEEE Vehicle Power and Propulsion Conference (VPPC), Anonymous pp. 1-6. [12] Logan, D., Pentzer, J., Brennan, S., 2012, "Comparing Batteries to Generators as Power Sources for use with Mobile Robotics," Journal of Power Sources, 212pp. 130-138. [13] M. M. Sadeghi, and E. F.

130-138. [13] M. M. Sadeghi, and E. F. Kececi, 2016, "Hybrid power system for mobile robotics," 2016 International Conference on System Reliability and Science (ICSRS), Anonymous pp. 64-67. [14] A. Sulaiman, F. Inambao, and G. Bright, 2013, "Development of solar hydrogen energy for mobile robots," 2013 6th Robotics and Mechatronics Conference (RobMech), Anonymous pp. 14- 19. [15] Amundson, K., Raade, J., Harding, N., 2006, "Development of Hybrid Hydraulic-Electric Power Units for Field and Service Robots," Advanced Robotics, 20(9) pp. 1015-1034. [16] Raade, J. W., and Kazerooni, H., 2005, "Analysis and Design of a Novel Hydraulic Power Source for Mobile Robots," IEEE Transactions on Automation Science and Engineering, 2(3) pp. 226-232. [17] Riofrio, J. A., and Barth, E. J., 2007, "A Free Piston Compressor as a Pneumatic Mobile Robot Power Supply: Design, Characterization and Experimental Operation," International Journal of Fluid Power, 8(1) pp. 17-28.

J., 2007, "A Free Piston Compressor as a Pneumatic Mobile Robot Power Supply: Design, Characterization and Experimental Operation," International Journal of Fluid Power, 8(1) pp. 17-28. [18] Sulaiman, A., Inambao, F., and Bright, G., 2014, "Solar-Hydrogen Energy as an Alternative Energy Source for Mobile Robots and the New-Age Car," IOP Conference Series: Materials Science and Engineering, 65. [19] Wilhelm, A. N., Surgenor, B. W., and Pharoah, J. G., 2006, "Design and Evaluation of a Micro-Fuel-Cell-Based Power System for a Mobile Robot," IEEE/ASME Transactions on Mechatronics, 11(4) pp. 471-476. [20] A. Mandow, J. M. Gomez-de-Gabriel, J. L. Martinez, 1996, "The Autonomous Mobile Robot AURORA for Greenhouse Operation," IEEE Robotics & Automation Magazine, 3(4) pp. 18-28. [21] D'Andrea, R., 2012, "A Revolution in the Warehouse: A Retrospective on Kiva Systems and the Grand Challenges Ahead," Ieee Transactions on Automation Science and Engineering, 9(4) pp. 638-639. [22] Eyerly, Jon V.

In general, the tensile, compressive and flexural strength are decreased against molded plastics. Some lasting products need the iteration of low flexural strength and high elongation to obtain an effective flexibility to bend in repetitive movements. The present work provides an analysis of the capacity of several selected thermoplastics materials such as Nylon (PA), Polyethylene Terephthalate (PETG), Polylactide (PLA), Polyurethane (TPU) and Polypropylene (PP) in order to test the maximal load capacity and the number of folding cycles sustained in perpendicular direction of movement. Results demonstrate that those of similar to injected molded products, PP and TPU materials surpass one million of cycles in the folding test.

Results demonstrate that those of similar to injected molded products, PP and TPU materials surpass one million of cycles in the folding test. Yet, in axial load they have lower strength against the other considered materials.Key Words: Folding Endurance; 3D Printing; Fusion Deposition ModelingIntroductionFlexible materials for Additive Manufacturing (AM) have been used by designers to develop novel textile structures [1] that can be converted in wearable products such as clothes, bracelets, neckless or shoes, while flexible mechanical parts can be printed in the shape of coil, leaf or torsion springs. 3D printed flexible electronics [2,3] are gaining terrain to produce flexible formats that are difficult or impossible to achieve with current technologies [4]. The field is so wide that new business models have been born to solve personalized needs to produce flexible orthoses [5].

The field is so wide that new business models have been born to solve personalized needs to produce flexible orthoses [5]. Living hinges are printed from simple plastic boxes, to functional articulations for prosthesis [6], robotic hands [7,8] or soft robots [9,10].FDM is one of the most used technologies in AM. Both, low cost FDM machines and the variety of filaments give new possibilities to create faster prototypes and new product development. Due to the opportunity of AM to print complicated shapes with no need of assemblies, the characterization of the FDM printed components is required.

Due to the opportunity of AM to print complicated shapes with no need of assemblies, the characterization of the FDM printed components is required. Many efforts to find mechanical properties have been restricted since the anisotropic behavior of the process; however, researchers in the area have evaluated mechanical tensile [11] and compressive strength [12], flexure force [13] and impact [14], but not the property to fold in repetitive cycles.2348Folding endurance is defined as the number of folds that are required to make a test piece break under standardized conditions [15]. Standard test methods are applicable to paper of thicknesses up to 0,25 mm thick [15,16]. There are no specific formulas that calculate a folding rate of plastics and no data sheets is proportioned by suppliers. Most commonly injected molding living hinges are made of PP, which can surpass 1 million of folding cycles [17] and a maximum load up to 20 Pounds/in [18].

Most commonly injected molding living hinges are made of PP, which can surpass 1 million of folding cycles [17] and a maximum load up to 20 Pounds/in [18]. Other materials used in folding injected molding products are: Polyvinyl Chloride (PVC), Urethane (UR) and PETG.ProcessFive 3D printing filaments were selected with different flexural strength and high elongation to obtain an effective flexibility from bending in repetitive movements: Nylon 12 (Stratasys, Ltd., Eden Prarie, MN, USA), PETG, PLA, PP (Gizmo Dorks, Temple City, CA, USA), and TPU Ninjaflex (NinjaTek, Manheim, PA, USA).An especial design for specimens was developed due that no standard test for plastic was appropriate for the objectives; transition heat factor provoked by the folding speed was not desirable, thus a folding endurance test for paper (ASTM-D2176-97a; ISO 5626:1993) was discarded. A living hinge design occurred to be the best option found to perform folding test.

A living hinge design occurred to be the best option found to perform folding test. First prototypes were development according with both the injected molding recommendations and the tool cooling living hinges [19] but the geometries and dimensions did not work for the FDM technology. Instead, applying all the recommendations from a design based on Smyth (2015) [20], a simple created design achieved favorable results (Fig. 1).Fig. 1 Living Hinge Specimen: Living hinges endurance was found by the accomplishment of the following two main tests:Test 1. Folding EnduranceA control station (Fig. 2) was built it with Arduino UNO (Arduino LLC) and 4 servomotors (JR ST47BB, Champaign, IL,USA) working at 2.22 times per second. Hinges were2349grasped with special holders and tighten with screws. The center of rotation of motor is aligned with the center of the folding area of the hinge.

Hinges were2349grasped with special holders and tighten with screws. The center of rotation of motor is aligned with the center of the folding area of the hinge. Controller was programmed to count the cycles of servomotors with 180 degrees of rotation, torque of 47 oz/in and automatic stops when the half of the hinge is fully detached. Nylon Specimens were printed on Stratasys Fortus 400mc (Stratasys, Ltd., Eden Prarie, MN, USA) and other materials in MakerBot Replicator Dual (MakerBot Industries, Brooklyn, NY, USA). Specimens were previously bent by 10 cycles at a temperature of 80oC to give orientation to the fibers according with the folding movement.ServomotorRotation ArmHinge HoldersLiving HingeFig. 2 Control Station for Folding Endurance TestTest 2 Static Axial LoadsSpecimens were attached in two specimen for fixtures dimensions (Fig. 3).

2 Control Station for Folding Endurance TestTest 2 Static Axial LoadsSpecimens were attached in two specimen for fixtures dimensions (Fig. 3). The Maximum Load Test consisted in measure the pressure with force gauge (Lutron FG-5100, Lutron Electronic Enterprise CO., LTD.) applied with a load cell (100 kgf) in axial direction in two different points listed below: 1. When the specimen starts to have a vertical movement below 100 microns. The goal is to find the maximum point when material starts its plastic deformation and is capable to carry weight with no affection in the precision of the folding movement. its 2. When specimen endures maximum the break point, resulting to obtain the maximum load that the material can support.fabricatedthethe load untilFig.

its 2. When specimen endures maximum the break point, resulting to obtain the maximum load that the material can support.fabricatedthethe load untilFig. 3 Static Axial Load2350Previous ExperimentBy avoiding the anisotropic behavior of the printed parts, an experiment was performed before the programmed tests to find the correct orientation of the printed specimens as well as to know the appropriated raster angle to get the maximum folding cycles. Following the criteria that orthogonal direction of the movement gets better strength, both XYZ and XZY [21] orientations were used. In the case of raster angles, a 0o and 45o/-45o were also considered. Fig. 4 represents viewpoints of the orientation of hinges and the direction of rasters. For rapid results PLA specimens were tested with printing parameters fixed deliberately at 0.25 mm of layer thickness with zero airgap, and two shells and 100% of infill.

For rapid results PLA specimens were tested with printing parameters fixed deliberately at 0.25 mm of layer thickness with zero airgap, and two shells and 100% of infill. PLA specimens were printed individually and not only heated at 80 Co but also folded 10 times before the experiment took place.Fig. 4 Representation of the Hinges and Initial Data of Previous ExperimentResultsThe number of cycles obtained in PLA for the previous experiment denoted no significant difference in the direction of the raster angle (0o or 45o/-45o), but was not the case to those of the 2 types of orientation (Fig. 5); particularly between the two factors used in the XYZ orientation, and the 45o/-45o raster angle used in XZY orientation. For subsequent tests, the XYZ orientation with the 0o raster angle with the highest mean (208.3 cycles) were utilized in the specimens volume of test criteria (1.5mm of length, 1mm of width and 25mm of height).2351Fig.

5 Plot of Folding Cycles in Previous ExperimentTest 1For the folding endurance test, PLA and Nylon specimens showed folding cycles below the 200 cycles in their means values,  (163.20 and 198.80 respectively) and PETG specimens broke before 900 cycles (=856.60, s=92.58). PP and TPU counted cycles resulted with an extraordinary differences among other materials. PP cycles raised more than 1 million whereas TPU test was stopped in 2 million, posteriorly, the specimens were physically checked and no damage in the surface of the folding area was observed. The results are tabulated in Table 1 and a Bonferroni differences are represented in Fig. 6. Despite that the number of cycles of PETG is larger in comparison with PLA and Nylon, no differences are denoted due to manly means of PP and TPU are very large.Table 1. Tabularized Folding Endurance Test.

Tabularized Folding Endurance Test. Results are in the number of cycles completed by different materialsSPECIMENPET (cycles)PLA (cycles)MATERIAL PP (cycles)TPU (cycles)NYLON (cycles)1 2 3 4 5733 920 960 876 794136 152 172 211 1451130002 980520 1107713 1080536 1060870+2000000 +2000000 +2000000 +2000000 +2000000192 126 170 276 230Mean Std Dev856.60 92.58163.20 29.831071928.20 57456.41+2000000 0.00198.80 57.252352Fig. 6 Bonferroni Differences Analysis in the Folding Endurance TestTest 2Axial loads were applied in perpendicular direction of the raster angle where most of the complete sets of hinges were broken in the area of test, but TPU, which in this case, printed rasters got broken in a different part of the hinge (Fig. 7).TPU SpecimenPLA SpecimenAxial Load LineFig.

7).TPU SpecimenPLA SpecimenAxial Load LineFig. 7 Differences in the Broken Line between TPU and PLA SpecimensTable 2 shows the data obtained in two points of the test: i) the applied force before a plastic deformation in the specimen and ii) the maximum load applied until the hinge is broken. Greater variability is observed in TPU and Nylon thermoplastics. In the case of TPU, results differ from overall mean that were observed (Fig. 8). With the exception of TPU material, large2353variability is presented in all samples tested, taking only 0.59 kgf to move TPU hinge in axial direction against the 6.86 kgf of the PET specimens.Table 2.

Results of Axial Load Test in Five Materials in Two Different Points of the TestSPECIMENMATERIALPET (kgf)PLA (kgf)PP (kgf)TPU (kgf)NYLON (kgf)Before DeformationMaximum LoadBefore DeformationMaximum LoadBefore DeformationMaximum LoadBefore DeformationMaximum LoadBefore DeformationMaximum Load1 2 3 4 58.60 8.55 5.30 6.30 5.5569.85 67.05 67.30 64.05 66.257.45 4.10 8.15 6.60 5.2070.25 83.60 81.30 82.80 85.153.75 5.30 3.60 5.20 5.6525.05 27.35 32.20 33.60 31.351.05 0.50 0.55 0.45 0.4035.40 24.00 19.55 30.10 27.259.40 4.20 7.40 7.45 4.3567.01 60.70 62.70 76.80 65.60Mean Std Dev6.86 1.6166.90 2.096.30 1.6580.62 5.964.70 0.9529.91 3.570.59 0.2627.26 6.016.56 2.2466.56 6.23Fig. 8 Interval Plot for the Response of Axial Force before any Deformation OccursInterval plot of the Maximum Load is presented in Fig. 9. Here the difference is notable between the hard and the elastic materials.

9. Here the difference is notable between the hard and the elastic materials. PP and TPU have no significant differences in the test before deformation process. Bonferroni simultaneous test confirms this as it is shown in Fig. 10, where an overlap can be seen with standard deviations of PET and Nylon. Also, PLA occurs to have the highest level of strength against axial load showing a statistic difference with PET and Nylon.2354Fig. 9 Interval Plot for the Response of Axial Force to a Maximum LoadFig. 10 Bonferroni Differences Analysis in the Axial Maximum LoadConclusionsThermoplastics materials used as filaments for fusion deposition modeling can be used in the manufacture of folding products for diverse applications.

PLA and Nylon materials result suitable for rapid prototyping of a few cycles folding objects, or for the rapid manufacture of non-lasting folding products like packaging, that support a good level to sustain adequately axial forces nearby 6 kg with no deformation. PETG occurs to work at 800 cycles above and with a 180o folding angular displacement, and an arithmetic mean that holds a static weight of 6.86 kgf, where the maximum load endurance mean resulted to be 66.90 kgf.2355PP semi-rigid material can behave as good as some folding injected molded products made by PP with an ability to be fold over 1 million of cycles. In the axial load test, PP was observed to be statistically similar to those of more rigid materials mentioned in above paragraph. TPU formulated thermoplastic (NinjaFlex Brand) was tested for 2 million of folding cycles with no visual damage.

TPU formulated thermoplastic (NinjaFlex Brand) was tested for 2 million of folding cycles with no visual damage. Yet, TPU could be used for applications with low strength to axial loads or designs with TPU inserts combined with other rigid materials.AcknowledgementsThe research presented here was conducted in the Rapid Prototyping Lab in the Autonomous University of Ciudad Juárez in Collaboration with The University of Texas at El Paso (UTEP) within the W.M. Keck Center for 3D Innovation.References[1] Melnikova, R., Ehrmann, A., & Finsterbusch, K. (2014). 3D printing of textile-based structures by Fused Deposition Modelling (FDM) with different polymer materials. In IOP Conference Series: Materials Science and Engineering (Vol. 62, No. 1, p. 012018). IOP Publishing.[2] Ko, S. H., Pan, H., Grigoropoulos, C. P., Luscombe, C. K., Fréchet, J. M., & Poulikakos, D. (2007). All-inkjet-printed flexible electronics fabrication on a polymer substrate metal by nanoparticles.

rced Convection in a Polymeric PowdersJeffery L. Norrell Graduate Research Asst.Kristin L. Wood Associate ProfessorRichard H. Crawford Associate ProfessorMechanical Engineering Department The University of Texas at AustinTheodore L. Bergman ProfessorMechanical Engineering Department The University of ConnecticutAbstractIn a Selective Laser Sintering (SLS) powder bed, thermal transfer occurs through multiple modes. Forced convection through the powder, or downdraft, has recently been implemented in SLS ma chines in an effort to enhance thermal transfer within the powder bed. In this paper, forced convec tion is analytically shown to be a significant thermal transfer mode for low porosities, such as seen in SLS powder beds. A polymeric powder bed subjected to downdraft is investigated with the goal of quantitatively determining thermal behavior. A numerical model describing heat transfer within a powder is presented.

A polymeric powder bed subjected to downdraft is investigated with the goal of quantitatively determining thermal behavior. A numerical model describing heat transfer within a powder is presented. The design and construction of an experimental apparatus to measure the temperature profiles within a powder subjected to forced convection is described.

Such a process able today. is inherently dominated by thermal processes, both between the laser and powder and within the powder bed.459Typical gas-saturated, polymeric powder beds complicate control of the thermal environment by being excellent thermal insulators due to the low thermal conductivity of both the polymer and saturating gas. Thermal control of the powder bed is critical to prevent unwanted effects such as part growth and warping. Many methods, such as radiative surface heaters and resistive wall heaters, have been attempted to achieve fine, rapid control over the thermal state of a powder bed. Another method, shown to have at least a qualitative effect on part quality, is forced convection through a powder. In current commercial SLS systems, a pressure drop is created below the powder bed to induce gas flow. Radiative heaters are used to heat the upper surface of the powder bed.

In current commercial SLS systems, a pressure drop is created below the powder bed to induce gas flow. Radiative heaters are used to heat the upper surface of the powder bed. In theory, gas then flows through this upper portion of the powder bed where the gas is heated. The gas then flows through the remaining portion of the powder bed, carrying thermal energy to the remainder of the powder.The goal of this work is to create a numerical model describing forced convection within apowder. Once such a model is obtained, experiments will be performed for model validation.Numerical ModelThe two-dimensional arrangements depicted in Figure 1, display the computational domains considered.

Figure l(a) is the domain for a pure conduction simulation, whereas Figure l(b) is the domain for a forced convection simulation, both with and without buoyant forces.f-- O.15m, 1j,I (6")-1I0.075m (3") ~H~Powder Bed, 1hxI I i-lInletI IO.3L (12") _t t<i(a)Powder BedOutlet0.25m (10") t1 J0.62m (24.5")(b)Figure 1: Computational domains.In this simulation, heated air, assumed to be slug flow, enters the top of the inlet. The heated air then passes through the powder bed, leaving the system via the outlet. An analytical model is created to describe heat transfer in this system. The governing equations for the pure fluid and the460saturated porous medium are given separately. that fluid flow is laminar, steady, incompressible, and two-dimensional.In writing the following equations, it is assumedFor the pure fluid region, the two-dimensional, steady state Navier-Stokes equations are used [4J.

av _ 0Conservation of mass is given by,au ox + oy -.Conservation of momentum within the pure fluid is described by,andp(u~: +v:;) = -~~ + ;x (JLf~:) + ;y (JLf~~) +pg~(T-Tc).Conservation of energy is expressed as,It is important to note that the thermophysical properties of the pure fluid are assumed constant. The exception is during buoyant simulations, where a Boussinesq approximation of fluid density, in which density varies with temperature, is used.For a fluid-saturated porous medium, the volumetrically-averaged laminar, inviscid, isothermalfluid flow is approximately described by the Darcy equation [1J,To account for non-isothermal fluid flow, a Brinkman-Forchheimer extended version of the Darcy model is used [2,5]. Within the saturated powder Darcian, or volumetrically-averaged, velocities are used.

Within the saturated powder Darcian, or volumetrically-averaged, velocities are used. Conservation of mass within the saturated powder is then given by,8UD &VD oX + oy == o.Conservation of momentum is given by,and,op o== - 8y + ox8 (8VD)OVD)0 (/-lef f ox + oy /-lef f oy + pgf3 (T - Tc) -(/-l1pC)!{ + VJ{IUDI VD·Equations (7) and (8) include the Brinkman extension (second and third terms) and the Forchheimer extension (the last term) to account for viscous and inertial effects, respectively.Conservation of energy is described by,Allthermophysical properties are again assumed constant, except for a Boussinesq density /-Lell == /-l1, which has beenapproximation in buoyant simulations.In Equations (7) and (8),461(1)(2)(3)(4)(5)(6)(7)(8)(9)previously seen to agree well with experimental data [8J.

It is important to note that Equations (6) through (8) neglect any radiation heat transfer between the fluid and powder or within the powder itself.The term kef f in Equation (9) is the effective thermal conductivity of the fluid-saturated powderbed. Though many expressions are available for kef f [6,7,9,IOJ, a simple weighted average,is used. The C term in equations (7) and (8) is the inertia coefficient given as [3J,C== 1.75 $-3/2. V175'The above analytical model is numerically solved for the two-dimensional, steady state tem perature distribution within both the fluid and fluid-saturated powder using the commercial finite element code FIDAP (version 7..5 from Fluid Dynamics International).ExperimentsTest MaterialsIn order to experimentally verify the numerical model presented above, an experimental ap paratus, shown in Figure 2(a), was constructed.

The working fluid for these experiments is air, while the powder bed consists of polycarbonate with a nominal particle diameter of 150 /-lm and an average poured porosity of 0.5. Regarding permeability, there are no general equations relat ing permeability as a function of porosity [5J. A number of empirical models exist, but all have restrictions in their application. If we consider equation (5), we see that for an isothermal system with known geometry, we can control flow rate, measure the pressure drop across the powder, and calculate a permeability. For a 0.025 m (I") thick, 0.25 m (I") diameter powder sample, subjected to a flow of 0.31 m/s (20 SCFH) to 0.62 m/s (40 SCFH), a pressure drop of approximately 1 kPa is measured. From Equation (5), we then determine that the powder permeability is approximately 1.5xl0-10 m2 .Experimental InstrumentationDimensions of the major sections of the experimental system are as shown in Figure 1(b).

In the experiment, compressed air passes over PID-controlled baffle heaters and then enters a plenum. Once through the plenum, the heated air passes through the inlet, through the powder bed, and leaves the system via the outlet.The baffle heaters are housed in a 0.10 m x 0.10 m (4" x 4") uninsulated plywood box. The baf fles are separated vertically by 0.00635 m (1/4"). The baffle box opens into aO.15 m x 0.15 m x 0.025 m (6" x 6" x I") plywood plenum, which sits atop the inlet. The inlet section is constructed of 0.15 m (6") ID polyvinyl chloride (PVC) pipe. Mounted on the outside of the inlet section is a 100 W PID controlled resistive heater covered with fiberglass insulation, used to maintain a quasi-isothermal boundary for the inlet. The inlet is mounted on top of the powder bed, which is housed in 0.15 m (6") ID steel pipe. The powder bed is surrounded by a fluid circulation jacket, also used to maintain an isothermal boundary.

The inlet is mounted on top of the powder bed, which is housed in 0.15 m (6") ID steel pipe. The powder bed is surrounded by a fluid circulation jacket, also used to maintain an isothermal boundary. The powder bed sits atop a fine wire mesh, allowing airflow through the powder. Pressure drop across the mesh alone was measured to be less than 1 Pa. Below the powder bed sits the outlet, constructed of 0.15 m (6") ID PVC pipe. Given the relatively low tempera tures in the outlet section and the high thermal mass of the PVC pipe, fiberglass insulation alone462(10)(11)#1=m/=l'll=m X -r('W) ~(1")()")0.025 m (1")#2#3#4XX#5#9#13#17#6X#10X#14 X#18 X#7X#11X#15 X#19 X#8 -+ #12 -+ #16 -+0.050m(2")X0.050m(2")X0.050 m (2")X0.050m (2")#20 --LX0.025 m (1")<LI(a) Apparatus.(b) Thermocouple numbering.Figure 2: Experimental system.sufficiently maintains an isothermal boundary condition.

An optional fan can be placed beneath the outlet to induce a pressure drop below the powder bed, further inciting airflow through the powder.Throughout the experimental device, 125 J-lm (5 mil), type K, PTFE-coated thermocouples are used to gather temperature data. A thermocouple is placed in the powder bed at each location marked w~th an "x" shown in Figure 2(b). Thermocouples also record air temperature in the inlet and outlet. All thermocouple data are collected and recorded with a Keithley 500A computer controlled data acquisition unit.Incoming compressed air can be heated from 25 0 C to 150 0 C, over a flow range of 0.017 mls (40 SCFH) to 0.034 mls (80 SCFH) in the inlet. The inlet wall can be heated from 25 0 C to 100 0 C.

The inlet wall can be heated from 25 0 C to 100 0 C. The fluid circulation jacket can be maintained over a range of temperatures from approximately 30 0 C to 100 0 C.The experimental system has several user-controlled inputs.Experimental ProcedureIn preparing for the experiment, the powder bed, wire mesh, and outlet, all initially at room temperature, are assembled together. Next, 0.025 m (I") of room temperature powder is poured inside. The steel powder bed wall is then repeatedly tapped, allowing the powder to settle. This procedure of pouring in powder then tapping is repeated, until the desired powder depth is achieved. The powder surface is then smoothed, and the remaining sections of the experimental device are assembled.Once initial assembly is completed, data acquisition is started. Airflow into the system and fluid flow into the fluid circulation jacket are then begun.

Airflow into the system and fluid flow into the fluid circulation jacket are then begun. The heated baffles and inlet wall heater463are next ramped to temperature over a period of 30 minutes.The system is left to run until thermal steady state is achieved. Typically, a run lasts from 8 to 15 hours. At this time, the recorded temperature data are downloaded and checked to verify that steady state has been achieved.ResultsTwo separate tests were run for comparison with the numerical model. For these two tests, the temperature set point for the incoming air was kept constant at 100 ° C, while the flow rate was varied. For each test, three numerical simulations were performed.The first test was run with an airflow of 0.017 mls (40 SCFH). Measured temperatures as a function of time are shown in Figure 3.

For each test, three numerical simulations were performed.The first test was run with an airflow of 0.017 mls (40 SCFH). Measured temperatures as a function of time are shown in Figure 3. Data curve labels indicate the thermocouple number, with Figure 2(b) indicating the in-bed position of the given thermocouple.For comparison, numerical simulations were also performed. Temperatures measured during the 40 SCFH experiment were used as boundary conditions in the simulations. Figure 4(a) shows the conduction heat transfer simulation for the powder bed. Figures 4(b) and 4(c) ,covering one-half of the experimental domain, show simulations that include both forced convection and conduction heat transfer. Figure 4(c), a buoyant simulation, also adds the Boussinesq air density approximation. The Ll T between isotherms in Figure 4 is 4° C.The second experiment was run with an airflow of 0.025 mls (60 SCFH).

The Ll T between isotherms in Figure 4 is 4° C.The second experiment was run with an airflow of 0.025 mls (60 SCFH). Figure 5 shows themeasured temperatures as a function of time.Figure 6 shows the numerical simulations run for the 60 SCFH experiment, where again Fig ure 6(a) is the conduction simulation, Figure 6(b) is the forced convection solution without buoyant forces, and Figure 6(c) is the convection solution with buoyant forces. As above, the Ll T ,between isotherms is 4 ° C.DiscussionIn terms of numerical results, by comparing Figure 4(a) to 4(b) and Figure 6(a) to 6(b) we see that, numerically, forced convection can have a significant impact on heat transfer within a polymeric powder bed. Keep in mind that Figures 4(a) and 6(a) represent the entire powder bed and Figures 4(b) and 6(b) represent half of the physical domain.

Keep in mind that Figures 4(a) and 6(a) represent the entire powder bed and Figures 4(b) and 6(b) represent half of the physical domain. Comparison is then made by considering isotherms of the same value in Figures 4(a), 4(b), and 4(c), and their location within the powder bed region. From this comparison, we can see that, in effect, forced convection works to transport thermal energy further into the powder.By considering Figure 4(b) versus 4(c) and Figure 6(b) versus 6(c), we can also see that natural convection, Le. buoyancy-induced flow, can have a significant effect at these low flow rates.

buoyancy-induced flow, can have a significant effect at these low flow rates. Also, from comparing Figure 4(b) to Figure 6(b), we see that increasing flow velocity through a powder increases thermal transfer into the deeper parts of the bed, as expected.In terms of the experimental results, we can see from both Figures 3 and 5 that achieving steady state, the point where dTI dt approaches zero and the temperature versus time curves begin to flatten out, can take a significant amount of time, over 7 hours. Air and polycarbonate, both relatively good thermal insulators alone, make an even better thermal insulator in the form of an air-saturated powder bed.If we compare the experimental data to the forced convection simulations, both with and without buoyancy, we begin to see a discrepancy.

118E+01 YMIIX FIDAP 7.52 21 Aug 96 11:36:19(c) Figure 6: Isotherms for 60 SCFH simulation.468..cantly lower than those predicted in the forced convection simulations. However, if the measured temperatures are compared to the conduction simulation, we see that there is excellent agreement. As in any research, a discrepancy between model and experiment implies a flaw in the model. In this instance, it is believed that the numerical model does not adequately reflect the air leaks present in the experimental apparatus. the resistance to flow imposed by the low permeability powder forces the air out of mating surfaces in the experiment. At present, these leaks prevent any experimentally driven conclusions about forced convection heat transfer within a polymeric powder.

At present, these leaks prevent any experimentally driven conclusions about forced convection heat transfer within a polymeric powder. However, we can easily see from Figures 3 and 5 that conduction heat transfer within a polycarbonate powder bed can take a significant amount of time to have any effect on the bed interior.In essence,Conclusions and Future WorkNumerically, forced convection heat transfer is seen to have a significant effect on steady state temperatures within an air-saturated polymeric powder bed. It is also numerically shown that, at sufficiently low flow rates, natural convection, or buoyancy-induced flow, must be accounted for in simulations.Experimentally, it is seen that the low permeability of fine powders, which are typically used in SLS, can have a drastic effect on attempts at generating flow through the powder. Much care must be taken to ensure that fluid is, in actuality, flowing through the powder.

Much care must be taken to ensure that fluid is, in actuality, flowing through the powder. Once true airflow through a powder is achieved and verified againstthe numerical model, fractional experiments will be performed. Control variables will be the initial air temperature, powder bed wall temperature, and airflow rate. In addition, a polymeric sintering model based on solid mechanics will be formulated and added to the numerical heat transfer. In conjunction with the model development, experiments will be undertaken wherein a heated object is embedded in the powder bed, reflecting the latent heat in a recently formed SLS part. In this way, sintering of the polymer under influence of forced convection can be studied. Information gained from these studies can then be used to better control the thermal state of a polymeric powder bed. More precise and rapid control may then help alleviate such nagging problems as growth and part warpage.Bibliography1.

More precise and rapid control may then help alleviate such nagging problems as growth and part warpage.Bibliography1. Bear, J., 1972, Dynamics of Fluids in Porous Media, American Elsevier, New York.2. Beckermann, C., Ramadhyani, S., and Viskanta, R., 1987, "Natural Convection Flow and Heat Transfer Between a Fluid Layer and a Porous Layer Inside a Rectangular Enclosure," J. Heat Trans., Vol. 109, pp. 363-370.3. Ergun, S., 1952, "Fluid Flow Through Packed Columns," Chem. Eng. Prog., Vol. 48, pp. 89-94.4. Incropera, F.P., and De Witt, D.P., 1990, Fundamentals of Heat and Mass Transfer, Third ed., John Wiley and Sons, New York, NY.5. Kaviany, M., 1991, Principles of Heat Transfer in Porous Media, Springer-Verlag, New York.6. Krupiczka, R., 1967, "Analysis of Thermal Conductivity in Granular Materials," J. Chem. Eng., Vol. 7, No.1, pp. 122-144.4697. Kubie, J., 1987, "Steady-State Conduction in Stagnant Beds of Solid Particles," Int. J. Heat and Mass Trans., Vol.

Chem. Eng., Vol. 7, No.1, pp. 122-144.4697. Kubie, J., 1987, "Steady-State Conduction in Stagnant Beds of Solid Particles," Int. J. Heat and Mass Trans., Vol. 30, No.5, pp. 937-947.8. Neale, G., and Nader, W., 1974, "Practical Significance of Brinkman's Extension of Darcy's Law: Coupled Parallel Flows Within a Channel and a Bounding Porous Medium," Canadian J. Chern. Eng., Vol. 52, pp. 475-478.9. Xue, S., and Barlow, J.W., 1991, "Models for the Prediction of the Thermal Conductivities of Powders," Proceedings, Solid Freeform Fabrication Symposium, H. Marcus, et al., ed., Austin, TX, pp. 62-69.10. Vagi, S., and Kunii, D., 1957, "Studies on Effective Thermal Conductivities in Packed Beds," J. AICHE, Vol. 3, No.3, pp. 373-381.470.

RENSIC APPLICATIONS OF SOLID FREEFORM FABRICATIONR.S. Crockett, R.ZickRapid Prototyping Center Milwaukee School of Engineering, Milwaukee, WI 53202AbstractSolid Freeform Fabrication was recently used to identify the dismembered body of a woman found in rural Wisconsin. Skin from the face of the victim had been removed, making visual identification impossible. A model of the skull was constructed by Laminated Object Manufacturing (LOM), using data extracted from CT scans of the victim’s head. Forensic anthropologists performed a facial reconstruction directly on the LOM model, which was then photographed and distributed. The computer model was further manipulated and served as supplemental data to investigators preparing the reconstruction. An identification from one of the distributed photographs led to the arrest of a suspect and a subsequent conviction.

An identification from one of the distributed photographs led to the arrest of a suspect and a subsequent conviction. This is the first example that we are aware of where SFF has been used in an active criminal case, and the successful results show a promising future for SFF as a forensic tool.BackgroundSolid Freeform Fabrication recently served an important role in a high-profile police investigation. The dismembered body of an unidentified woman was found in the Wisconsin River near Spring Green, Wisconsin in July, 1999. Skin from the face of the victim had been removed, thus making visual identification impossible. Although an autopsy revealed information about the victim’s sex, age, race, and medical history, investigators could not find a match with any reported missing person in national databases.

Although an autopsy revealed information about the victim’s sex, age, race, and medical history, investigators could not find a match with any reported missing person in national databases. With no further leads, it was decided that a visual identification from a facial reconstruction was the most promising approach.Forensic anthropologists routinely create facial reconstructions from discovered skulls, whereby clay is built up over multiple reference points on the skull surface. The height of these points corresponds to statistically determined tissue depth in various areas (Figure 1). In this particular case, however, the standard procedure for a facial reconstruction was not a desirable option for identifying the victim. Because the victim’s remains were an actual head, not merely bone, the remaining flesh would have had to be removed to prepare the skull for reconstruction.

Because the victim’s remains were an actual head, not merely bone, the remaining flesh would have had to be removed to prepare the skull for reconstruction. This remaining soft tissue needed to be preserved as possible evidence, as it contained cut marks and other identifying clues to the murder.549Figure 1. Typical forensic reconstruction. Pegs are rubber erasers, cut to lengths determined by statistical data.Milwaukee School of Engineering received a call from Detective Joe Welsch of the Sauk County Sheriff’s Department, indicating that he had been informed of the capabilities of Solid Freeform Fabrication and asking if we could help.Model PreparationThe victim’s head was CT imaged at the University of Wisconsin, and the resulting data file was transmitted to the Rapid Prototyping Center at MSOE. Mimics software from Materialise1 was used to extract just the features of the skull (Figure 2). During this process, minor CAD manipulation of the data file was required.

Mimics software from Materialise1 was used to extract just the features of the skull (Figure 2). During this process, minor CAD manipulation of the data file was required. The top of the skull had been removed during the autopsy; although this cut was fairly clean in the resulting computer solid model, Mimics tools were used on the model to ensure that it was a separate piece. Similar processing was performed on the jaw. In addition, it was decided to fill the interior of the skull model with solid material, both to speed up the LOM buildFigure 2. (left) Mimics software was used to manipulate the CT data and extract just the features of the victim’s skull. (right) Completed computer model1 http://www.materialise.com550process and to give the resulting physical model a density and center of gravity similar to an actual skull. Finally, a CAD-designed attachment was merged with the model to enable the SFF object to be supported on a stand during the reconstruction.

Finally, a CAD-designed attachment was merged with the model to enable the SFF object to be supported on a stand during the reconstruction. Anthropologists typically make use of the opening at the base of a skull where the spinal cord terminates. Because we were building a filled skull, however, we had to replicate this feature artificially. Later feedback from the anthropologist confirmed that adding designed components such as this to an otherwise organic structure could significantly simplify their process.An .stl file was generated directly from Mimics, and a full-scale, accurate model of the skull was then build using Laminated Object Manufacturing (Figure 3). Although MSOE also has equipment for Stereolithography, Fused Deposition Modeling, and Selective Laser Sintering, LOM was preferred both because it was the fastest machine for this size and geometry (filled solid with a complex surface), and because the resulting physical model has a density and “warmth” similar to bone.

This second fact was extremely important, as it enabled the acceptance of SFF technology by extremely conservative professions such as law enforcement and anthropology, as well as by a later jury. Total build time on the LOM was 30 hours. Within 3 days from scan data to physical object, the skull model was shipped to a forensic anthropologist from the Kentucky State Medical Examiner’s Office.Computer Model Data ManipulationAlthough we did not plan it at the time, the .stl file used to create the skull model served an important continuing role as a source of supplemental data during the process of creating the facial reconstruction.

First, at the request of the anthropologist, features of soft tissue still extant on the skull were manipulated on the computer directly from CT data; measurements and printouts of the skull both with and without this tissue were used to aid in the forensic reconstruction (Figure 4).Additionally, measurements taken directly from the .stl file that created the skull wereFigure 3. (left) LOM model before decubing. (right) Finished model.551Figure 4. Computer model was further manipulated and used as supplemental data to aid the forensic anthropologist. (left) Imaging of the soft tissue. (right) Measurements used for confirmation of victim’s race.used to confirm the victim’s race. At the completion of the facial reconstruction, many of the resulting features appeared to be Asian, whereas the earlier autopsy had determined the race to be black.

At the completion of the facial reconstruction, many of the resulting features appeared to be Asian, whereas the earlier autopsy had determined the race to be black. Forensic reconstructions are a combination of science, art, and the intuition of the anthropologist, so this new information was taken very seriously. A typical method of determining race involves taking measurements of the skull at about 30 points using calipers; statistically, these measurements have a high correlation with race. Unfortunately, by the time the question of race was raised, the reconstruction was nearly complete. To take measurements on the physical model would require removing the clay; to take measurements on the actual head would require removing the remaining soft tissue, which was not an option. The solution was to take measurements directly off the .stl file, using point-point measurement tools in Magics RP (also from Materialise).

The solution was to take measurements directly off the .stl file, using point-point measurement tools in Magics RP (also from Materialise). These measurements confirmed the race of the victim, again avoiding any disturbance to the actual evidence.Facial Reconstruction and ResultsThe completed reconstruction was photographed with a variety of hairstyles. Investigators selected four images and prepared flyers, which were posted around the area of the murder (Figure 5). Nearly three months later, someone identified one of the images. The woman was a student from Tanzania who had not been reported missing.552Figure 5. (left) Photographs of the completed reconstruction were distributed by the Sheriff’s department. (right) An identification of the victim was made by the above photograph, which was described by the identifier as an exact match to the victim.According to investigators, the image (shown enlarged in Figure 5) was nearly identical to the victim.

Within a day, police had arrested her cousin and charged him with first degree homicide. He was subsequently convicted and is currently serving a life sentence. Without the identification of the victim, it is likely that this murder would have gone unsolved.The Future of SFF as a Forensic ToolThis was a highly visible and successful technology demonstration, and it points to some promising continued use of SFF as a forensic tool. The FBI and the National Institute of Justice have both expressed interest in this technique as a near-term solution to automating the identification of victims.

The FBI and the National Institute of Justice have both expressed interest in this technique as a near-term solution to automating the identification of victims. There are three primary areas where SFF could potentially reduce the time and cost, as well as increase the consistency, of facial reconstructions:1) Integration of CAD reconstruction.1) Integration of CAD reconstruction.intothe processto reducethe preparationtimeA simple example of this would be the tissue depth markers of Figure 1 placed during feature extraction and built directly into the model. The ultimate goal would be to deliver a skull model to the anthropologist that is nearly complete, with base tissue depth already part of the physical model.5532) Ability to extract features of interest and manipulate data in physical form.It has been shown that features of the bone can be selectively extracted from a larger data set.

Alternatively, SFF models of just the soft tissue can be constructed, as well as interior sections, etc. All of this data in physical form is important to anthropologists, who are used to working with physical objects.3) Ability to create multiple copiesSFF can be used to create multiple copies of forensic evidence, for multiple reconstructions or multiple partial reconstructions.This is the first example that we are aware of where SFF has been used in an active criminal case. All of the participants in this project were really stretching the limits of what can be done – not from a technology point of view, but from an emotional or mental point of view. New tools such as DNA sequencing in the 1980s and Solid Freeform Fabrication today can make a huge impact on the way crimes are solved. Acceptance of these innovations, however, is generally much slower.

Acceptance of these innovations, however, is generally much slower. By using SFF in a successful case, an important step has been taken toward making this technique available to a larger community.ACKNOWLEDGEMENTSThe authors would like to acknowledge the Sauk County Sheriffs Department, especially Detective Joe Welsh; Leslie Eisenberg of the State Historical Society of Wisconsin, who first suggested SFF as a solution; Dr. Emily Craig of the Kentucky State Medical Examiner’s Office, who performed the reconstruction; and Dr. Lisa Milkowski of MSOE, who assisted in the data manipulation. We also gratefully acknowledge the financial support of the National Science Foundation and the MSOE Rapid Prototyping Consortium, as well as the facilities and expertise provided by the staff of the MSOE Rapid Prototyping Center.554.

RM ACCURACY ANALYSIS OF CYLINDRICAL PARTS PRODUCED BY RAPID PROTOTYPING G. R. N. Tagore*, Swapnil. D. Anjikar+, A. Venu Gopal* * Department of Mechanical Engineering, National Institute of Technology, Warangal + Research Scholar, Dept. of Mechanical Engineering, NIT–Warangal INDIA.AbstractSolid Freeform fabrication processes are being considered for creating fit and assembly nature functional parts. It is extremely important that these parts are within allowable dimensional and geometric tolerance. The part accuracy produced by rapid prototyping process is greatly affected by the relative orientation of build and face normal directions. A systematic method is needed to find the reliability of the created product. This paper discusses the work done in this area and the effect of build orientation on the part form accuracy analysis of each specified tolerance like circularity and cylindricity. Feasible build direction that can be used to satisfy those tolerances is identified.

Feasible build direction that can be used to satisfy those tolerances is identified. It will help process engineer in selecting a build direction that can satisfy a mathematical model of form tolerance.1. INTRODUCTIONThe main advantages of Solid Freeform Fabrication (SFF) processes are that they don’t require any part specific tooling and completely automated. In Fused Deposition Modeling (FDM) a three dimensional CAD model of the part is sliced into layers and the numerical data on the geometry of layers is then feed into the fabrication unit, which builds each layer sequentially until the entire part is fabricated.SFF processes are conducive to the concept of distributed design and manufacturing where in process providers will list their constraint on website and designers will perform manufacturability analysis for their design. This helps to reduce redesign when manufacturing constraint are violated.

This helps to reduce redesign when manufacturing constraint are violated. In order to create defect free functional parts through rapid prototyping techniques, manufacturability of design tolerance with respect to process constraints is essential. Until recently SFF processes were primarily used for creating prototype parts. Increasingly SFF processes are being considered for creating functional parts. In such applications, SFF can either be used for creating tooling i.e., patterns for casting, low volume molds, etc. or directly creating the functional part itself. In order to create defect free functional parts, it is extremely important to fabricate the parts within allowable dimensional and geometric tolerances. In order to determine whether a process can produce the part within required tolerances, we need to analyze manufacturability of design tolerances with respect to process constraints. Industries use 15.6% of parts produced by SFF by fit and assembly tests.

Industries use 15.6% of parts produced by SFF by fit and assembly tests. Fit and assembly tests need evaluation of form errors for attaining correct fit during assembly. In this paper we primarily focus on circularity and cylindricity tolerances assigned to the cylindrical parts.Fused deposition modeling (FDM) is a rapid prototyping (RP) process that fabricates parts layer by layer by deposition of molten thermoplastic material extruded from a nozzle. A proprietary software, Quickslice, processes the STL file to create the slices and roads and commands the FDM machine to generate layers of specified thickness and road width from the nozzle of a liquefier head. In general, the outer perimeter of the layer is laid down first, after which fill roads are created to fill the solid areas inside each layer. The types of fill patterns available are the raster, the contour or a combination of both.

The types of fill patterns available are the raster, the contour or a combination of both. The layers are depositedcontinuously at any part build orientation to build the part bottom up. Geometric accuracy of components is one of the most important quality characteristics in layered manufacturing processes on which most rapid prototyping (RP) techniques are based. Layered manufacturing is an approximate fabricating process in which the final geometric error of the physical part is affected, not only by the approximation technique used, but also by the fabrication process.2. ERRORS INDUCED IN LAYERED MANUFACTURINGAlthough the original RPTs were developed to provide style models, current RPTs are increasingly regarded as manufacturing technologies for tooling and production parts [3]. All manufacturing technologies for tooling and production parts must guarantee the accuracy of geometric dimensions and repeatability.

All manufacturing technologies for tooling and production parts must guarantee the accuracy of geometric dimensions and repeatability. The subject of geometric accuracy of parts is thus under intense debate as practitioners argue where and under what circumstances a rapid prototyping machine provides enough accuracy to be suitable for making tools [4]. Yan and Gu [5] categorised the most common sources of errors in rapid prototyping processes into mathematical error, process-related error and material-related error. Chalasani and Bagchi [6] identified tessellation, slicing orientation and slicing location as the three primary sources of error in the data preparation stage. Also, errors in the data preparation process were found in facet approximation, staircase effect and containment problem, and in laser beam path planning algorithms [7]. Process-related errors affect the shape of the layer in the x,y-plane and along the z-axis as well as the overall 3D shape [8].

Process-related errors affect the shape of the layer in the x,y-plane and along the z-axis as well as the overall 3D shape [8]. The physical and chemical characteristics of the materials in the process of solidification involving liquid and powder materials have been studied by many workers [9] who showed that the principal process-related errors are shrinkage and distortion in using these materials. For solid sheet materials, little is reported for analysing component geometrical error relating to the fabrication process. From the viewpoint of component production process in layered manufacturing, it is very important to understand the error distribution and error transfer mechanisms in order to assess, predict and control the final geometric accuracy of the part.

A geometrical model is used for assessing the effects of various errors and their accumulation and the final part geometry in RPT considering error interactions between the data preparation process and fabrication process.2.1 Data Preparation Error (DPE)In the RPT process, each 2½-D layer is generated as a "sweep" of a planar profile by depositing material on the interior of a 2D slice. Since the whole part is manufactured in this manner, the boundary of the part created is a stepped approximation of the boundary of the ideal part. As a result of this, compared with the parts created in the conventional manufacturing technologies such as milling and turning, parts produced in layered manufacturing exhibit a staircase effect. The error resulting from the staircase effect can be affected by the containment mode which describes the geometrical arrangement between the ideal part geometry and the sliced part contours[1].

The error resulting from the staircase effect can be affected by the containment mode which describes the geometrical arrangement between the ideal part geometry and the sliced part contours[1]. There are three containment modes to approximate the ideal part in the RPT process. These are referred to as the positive situation (the sliced part profile slightly exceeding the ideal profile), negative situation (the sliced profiles situated within the ideal profile) and the hybrid situation [7]. The approximating accuracy in the slicing procedure also depends on the slicing orientation, location and slicing thickness as reported in [6]. These errors are not related to the actual fabricating process and on1y depend on the fabricating method selected and slicing parameters used.

These errors are not related to the actual fabricating process and on1y depend on the fabricating method selected and slicing parameters used. These errors are therefore lumped as data preparation error(DPE) and they can be predicted theoretically by analyzing the extent of the staircase effect for a given containment situation and slicing parameters selected.2.2 Fabrication Process ErrorIn the actual RPT part-building process there are additional errors which are affected by many process-related parameters, e.g. the machine path control accuracy, tool scan speed uniformity, tool shape stability, platform control accuracy, material properties, part thermal distortion, material feed uniformity, fixture stability, glue thickness (in LOM process) uniformity, and part thermal shrinkage and distortion. These errors interact with the DPE and when combined they determine the final geometrical accuracy of the part produced.

These errors interact with the DPE and when combined they determine the final geometrical accuracy of the part produced. There are three possible effects of these fabrication errors on the final geometry of the part:1. Geometrical errors generated in the x- and y-directions only which are primarily associated with the scan path and tool shape control accuracy as shown in Fig. l(a).2. Geometrical errors in the z-direction which are mainly related to the platform displacement accuracy and materials thickness uniformity (e.g. in LOM and SLS processes) as shown in Fig. l(b).3. Geometrical errors in all directions (Fig. l(c)) as a result of scan speed variation (acceleration and deceleration), material flatness variation (e.g. in LOM process), thermal distortion of the part and other assembly errors.

l(c)) as a result of scan speed variation (acceleration and deceleration), material flatness variation (e.g. in LOM process), thermal distortion of the part and other assembly errors. It can be seen that the errors resulting from the actual part building process can alter the extent of theoretical errors relating to the staircase and containment problem defined by the fabrication method. For the purpose of error analysis, all the errors generated during the part-building process are now defined as the disturbance error.Fig.1. Disturbances in the RP part building process (a) Errors in the x,y-plane (e.g. contour geometry error). (b) Errors in the z-direction (e.g. layer thickness error). (c) Errors in x-, y-, and z-directions (e.g. layer flatness variation)3. METHODOLOGYGeometric dimensioning achieves the goal of identical dimensions through four simpleobvious steps.1. Convey the nominal distances and orientations from origin to other surfaces. 2.

METHODOLOGYGeometric dimensioning achieves the goal of identical dimensions through four simpleobvious steps.1. Convey the nominal distances and orientations from origin to other surfaces. 2. Establish boundaries / or tolerance zone for specific attributes of each surface along with specific rules for conformance3. Allow dynamic interaction between tolerances (Simulating actual assembly possibilities) where appropriates to maximize tolerances.A form tolerance is applied to a non-datum feature only where there is some risk that the surface will be manufactured which form deviations giving enough to cause problems in sub sequent manufacturing operations, inspection, assembly or function of the part. Circularity is a condition of surface. Circularity tolerance is a refinement of the size tolerance.

Circularity is a condition of surface. Circularity tolerance is a refinement of the size tolerance. In case of cylinder each circular element is independent of every other circular element that means that the part can look like stack of pennies that is misaligned but that can still satisfy a circularity inspection [10-11]. A circularity tolerance controls feature’s circularity (roundness) at individual cross section. While sweeping the tolerance zone may continually adjust in overall size, but shall maintain the specified radial width. This effectively removes diametrical taper circularity control. Additionally this spine orientation and curvature may be adjusted within the constraints mentioned already. This effectively removes axial straightness from circularity control. The circularity tolerance zone need not be concentric with either size limit boundary.In a layer manufacturing process such as FDM the part surface geometry is approximated along the building direction.

The curved features are formed at the edges of a two adjacent slices. This results in an error resulting a staircase effect. The actual geometrical formed may be different, depending on the geometry of the part and orientation of part in the building process. It is assumed that the errors don’t transfer between layer during the data preparation phase. The fabrication process will transfer the errors only in the part building direction.Cylindricity is a condition where all points on the surface of a cylinder are equidistance from the axes. Unlike circularity the cylindricity tolerance applies to circular and longitudinal elements at the same time[11]. Cylindricity is a composite form tolerance that simultaneously controls circularity, straightness of a surface, and taper of cylindrical features. A roller bearing might be controlled with a cylindricity tolerance where as a conical surface (bearing) might be better controlled with profile tolerance.

A roller bearing might be controlled with a cylindricity tolerance where as a conical surface (bearing) might be better controlled with profile tolerance. When the size tolerance doesn’t control the form of a feature, a form tolerance may be specified as a refinement to cylindrical tolerance zones or material conditions which are appropriate for surface control.It specifies a tolerance zone bounded by two concentric cylinders whose Radii differ by an amount equal to the tolerance value. The tolerance zone cylinder may adjust to any diameter, provided their radial separation remains equal to the tolerance value. This effectively removes feature size from the cylindricity control.Co-axiality is the relationship between multiple cylindrical or revolute features sharing a common axis. Coaxiality can be specified using run out, concentricity, or positional tolerance. A runout tolerance controls surface deviations directly, without regard for the features axes.

Coaxiality can be specified using run out, concentricity, or positional tolerance. A runout tolerance controls surface deviations directly, without regard for the features axes. A concentricity tolerance controls the mid points of the diametrically opposite points. Coplanarity can be specified using either a symmetry or positional tolerance. A symmetry tolerance controls the mid points of opposed surface points.Runout is the variation in the surface elements of round feature relative to an axis. In precision assembly, run out causes misalignment and or balance problems [11]. Examples are run out of the ring groove diameters relative to the pistons diameters might cause the ring to squeeze unevenly around the piston or force the piston off centre in its goal. A motor shaft that run out relative to its bearings will cause the motors to run out of balance, shortening its working life. It is called wobble and lop sidedness.

A motor shaft that run out relative to its bearings will cause the motors to run out of balance, shortening its working life. It is called wobble and lop sidedness. Total run out adds further refinements to the requirements of circular run outs. FIM (full indicator movement) is the difference in mm between an indicator most positive and most negative movements during run out. It can be measured using CMM. Circular run out tolerance applies at every possible circle on the feature’ssurface but each circle may be evaluated separately from the others. One should add run out tolerance for each journal to refine the feature’s tolerances.If the FIM reading difference for each circle is below the run out tolerance, it can be said that the part satisfies the run out tolerances. In case total run out, its tolerance sweeps over the entire control surface rather than each circular element being evaluated separately.

In case total run out, its tolerance sweeps over the entire control surface rather than each circular element being evaluated separately. The total run out FIM encompasses the highest and lowest of all readings obtained at all circles. Any taper in the controlled feature will increase the FIM.Circular run out tolerance is often ideal for ‘O’ ring groove diameters. Run out tolerancesare especially suited where alignments and dynamic balances are critical.Circular run out = circularity + concentricity Total run out = cylindricity + concentricity Runout tolerance applies directly to surface elements where as positional tolerance RFS controls only the coaxiality of feature’s actual mating envelop. It doesn’t provide form control for the surface. Position tolerance is considered instead of run out when interaction is desirable and size limits will adequately control form. A feature’s run out tolerance need not be less than size tolerance.

Position tolerance is considered instead of run out when interaction is desirable and size limits will adequately control form. A feature’s run out tolerance need not be less than size tolerance. Hence each run out tolerance has to be considered independently and carefully. A few well thought out run out tolerances are evolved to control combinations of relationships. Run out tolerance yields a worst case inner boundary equal in size to the feature’s small limits size–the value of its run out tolerance and a worst case outer boundary equal in size to the feature’s large limit size + the value of its run out tolerance. The inner or outer boundary can be exploited to protect a secondary requirement for clearance without using a separate positional tolerance.4. MATHEMATICAL MODELThe entire feature surface shall every where be contained within the tolerance zone. The interpretation of semantics of tolerance is burning topic in form tolerance.

MATHEMATICAL MODELThe entire feature surface shall every where be contained within the tolerance zone. The interpretation of semantics of tolerance is burning topic in form tolerance. The variance area of each tolerance zone can be derived. The tolerance zone means the region or area which limits the real feature’s movement. The real feature is feature with error, and is composed of points in the nominal feature after some ways of movement. Tolerance of a cylindrical feature is concerned with the constraint on the cylinder surface or on the cylindrical axes. For the same feature, the matrix equation of different tolerances can be the same. But the movement ranges of degree variables in the equations are generally different. Hence, establishment of the constraint model of each degree variable according to tolerance definition is the urge of the day. In this work each tolerance zone’ mathematical model is established with inequality based on the degree of the feature.

In this work each tolerance zone’ mathematical model is established with inequality based on the degree of the feature. Each of the symbols is as follows:Cd is the direction vector of tolerance zone; Cp is the position vector of tolerance zone; P is a point in the tolerance zone; r is the normal radius of the cylinder; es, ei are the upper and lower specification limits of the cylinder’ diameter; Tcir is the form tolerance; dx, dy, dz are theFig 2 Definition of Circularity tolerance zone Fig 3. Maximum shift of circularity toleranceFig.4 Cylindricity tolerance zone definition of cylindricity tolerance zoneFig 5 Maximum orientation anglez represents translational along the axial and rotation range around the axial is δθ,δφ and δψ respectively. Generally a complete tolerance specification is constructed with the default tolerance specification and the specified tolerance.

Generally a complete tolerance specification is constructed with the default tolerance specification and the specified tolerance. Here the correctness cannot be guaranteed for such tolerance specification and following three aspects should be considered:translation range of the coordinate origin along the axes x, y and z respectively. δ , δ δ1. Is the given tolerance and size reasonable? 2. Is the given tolerance complete? 3. Is the given tolerance specification valid? The tolerance validity is evaluated by generating variational geometries. The tolerance semantics has four basic attributes say position, direction, form and size. The size is specified by the designers. Hence, the key to exactly represent tolerance semantics is to determine the position and direction of the tolerance zone. The traditional tolerance types like circularity and cylindricity do not convey the real meaning. Hence, tolerance zone floated is new way of classification of tolerance.

The traditional tolerance types like circularity and cylindricity do not convey the real meaning. Hence, tolerance zone floated is new way of classification of tolerance. The variational geometry can be generated by generating all the variational elements like surface and axis and then regenerate the whole variation part through the surface to surface or curve to curve or surface to curve intersection. For a cylinder surface it is sliced equally into many equal segments along the direction perpendicular to the direction. The variational cylinder surface of the each slice is generated which satisfies the tolerance requirement. Then skinning operation of the generated surface completes the process.xy,The study helps integration of CAD/CAM and interpreting the semantics of form tolerances exactly and completely.

Then skinning operation of the generated surface completes the process.xy,The study helps integration of CAD/CAM and interpreting the semantics of form tolerances exactly and completely. It helps the designer to fix the correct cylindrical tolerances to real production experience that facilitates assemblies in groups that meet the functional requirements based on variance area of each tolerance zone [12].The mathematical model of the hole tolerance is as follows:The inequality equation of each degree variables’ movement range of cylindricitytolerance is as follows.−Tcy2≤(y+δ y+x . δ ψ)2+(z+δ z+x . δ φ)2r ≤−Tcy2……….1WhereTcy / 2 < δy < Tcy / 2, - Tcy / 2 < δz < Tcy / 2……….2The inequality in equation 1 shows points in the cylindricity tolerance zone anddetermines the size and form of the cylindricity tolerance zone.

allowable deviation0.079011 0.072514 0.06526 0.071253 0.061501 0.053996 0.064498 0.070251 0.065257 0.059999 0.057754 0.084738 0.063748 0.078769 0.079011The consistency of the data point mentioned above is tested using Chauvenet’s criterion and it is found that there is no inconsistency in the test data points. From these reading it is observed that the circularity, diametrical and runout errors are minimum for zero degree orientation for the cylinder. The cylindricity error is also minimum at zero degree orientation. However the maximum translation error is more at zero degree orientation. This clearly shows that decisions cannot be based purely on conventional form error like circularity error, cylindrcity and dimensional error. The translational error and rotational error are much more important in assembly operations and dynamic service condition.

The translational error and rotational error are much more important in assembly operations and dynamic service condition. Hence a tolerance zone mathematical model established by inequality based on the degree of feature is more important to take care of translational errors.6. CONCLUSION 1. Real feature can be called feature with error due to translation and rotation and is composed of points in the nominal feature after some ways of movement. Hence a generic mathematical model is developed which can use real feature tolerance for calculating the form error to give more accurate results. This approach will reduce the difficulties in assembly operation. 2. A tolerance zone mathematical model is established and variance area of each tolerance zone is obtained. The equations developed can be used for determining the maximum transfer along the axis. 3. This approach helps to create simulation cylinder of the hole during assembly or dynamic operation.

The equations developed can be used for determining the maximum transfer along the axis. 3. This approach helps to create simulation cylinder of the hole during assembly or dynamic operation. This simulation model will help in identification of the possibilities of seizure of piston during its operation in cylinder or seizure of bearings for shaft.4. This tolerance model is essential for integration of CAD and CAM. This integration can mitigate the problems while bridging the gap between CAD and CAMACKNOWLEDGEMENT The Authors wish to express their gratitude to The Director, National Institute of Technology Warangal India, for permitting us to carry out the work and also permitting us to present the paper at 18th SFF International Symposium at University of Texas Austin.REFERENCES1. W. Liu, L. Li and A. K.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFormation of Easy-To-Remove Supports in Laser Powder Bed Fusion through Selective DopingM. G. Sperry*, D. Carter*, N. B. Crane*, T. W. Nelson*Ira A. Fulton College of Engineering, Brigham Young University, Provo, UT 84604AbstractLaser Powder Bed Fusion (LPBF) is a popular Additive Manufacturing (AM) technique used c01m11only for metals. Metal paiis formed by LPBF generally require supports connecting the part to the print bed to hold up the strncture, remove heat, and minimize deformation due to solidification shrinkage. Because of these suppo1is, finished parts must be cut away from the build plate, and generally require additional machining to achieve the desired geometry. In this study, a carbon suspension was deposited in the 3 l 6L stainless steel powder bed at the interface between the support and the finished part.

In this study, a carbon suspension was deposited in the 3 l 6L stainless steel powder bed at the interface between the support and the finished part. The added carbon reduces the co1Tosion resistance of the 3 l 6L. This allows full fusion of the support material to provide heat transfer and mechanical support during printing, while allowing the supports to etch preferentially by electrolytic etching. This causes the finished part to etch or break free from the supports without any need for machining, simplifying post-processing.IntroductionLaser Powder Bed Fusion (LPBF) is an Additive Manufacturing (AM) technique that forms 3-dimensional designs by selectively fusing powder one layer at a time with a laser [ l ]. The process allows for internal voids, curves, and other geometries that are costly or impossible to produce with traditional casting and machining processes.

The process allows for internal voids, curves, and other geometries that are costly or impossible to produce with traditional casting and machining processes. This gives designers flexibility to produce parts with unprecedented complexity, and customizability [2].In LPBF of metal, paiis require sufficient support not only to hold up overhanging features, but also to resist sluinkage deformation [3]. They also provide an critical heat conduit to cool the melt pools [ l ]. The geometry and placement of supports is a critical aspect of the AM process [ 4, 5]. In LPBF, supports are welded to both the build plate and the finished part, making support removal a multi-step process. For robust support systems, as shown in Figure 1, parts are generally heat treated to relieve stress, then cut away from the build plate with a band saw or wire EDM and machined with a mill or ground by hand to remove the remaining supports.

Multi-step post processing of metal LPBF parts reduces efficiency and may require over-design with sacrificial material to allow for support removal [6, 7]. Polymer PBF often avoids defom1ation by maintaining the bed above the crystallization temperature [8]. Vora, et al. showed how a similar benefit468could be obtained in metal LPBF using in situ alloying to locally create a eutectic composition [9]...Figure l. Complex design formed with Steel by LPBF. Extensivesuppo1i structures such as these require multi-step post processing.Meanwhile, some AM processes, such as plastic filament printing allow for the use of soluble support material. These dissolvable suppo1is allow for delicate and intricate features, including cantilevers, horizontal bores, and long spans without concern for damage during support removal [ l 0].

These dissolvable suppo1is allow for delicate and intricate features, including cantilevers, horizontal bores, and long spans without concern for damage during support removal [ l 0]. Dissolvable supports decrease the fabrication time and cost as post processing is simplified by removing machining requirements [11] [12].Wei et al showed that similar benefits could be obtained in LPBF by using a weak secondary material for support. However, this method requires a set of high precision vacuum and deposition nozzles to selectively remove material from the print bed, then deposit a secondary material where supp01is will be forn1ed [5]. This method adds significant time to the print process and reduces the heat transfer capacity of the supp01is [5]. Other metal support removal methods proposed by Lefky et al. and Hildreth et al. require the addition of a sacrificial shell that is uniformly corroded away from the entire part [7, 13].

Other metal support removal methods proposed by Lefky et al. and Hildreth et al. require the addition of a sacrificial shell that is uniformly corroded away from the entire part [7, 13]. These methods reduce the resolution of negative features attainable by LPBF and increase the number of post-processing steps by adding a carbmizing period.This paper explores the effects of adding carbon as a corrosion-inducing agent to support material in a stainless steel (grade 3 l 6L) powder bed, then applying an electrolytic etch that has minimal influence on SS 316L but causes significant co1Tosion of the modified mate1ial.

Electrochemical machining (ECM) takes advantage of the chemistry and electrical potentials of a work piece to precisely add or remove material using plating or etching fluids [ 14, 15]; ECM and its variants are the topics of hundreds of research papers annually [ 16], which provides excellent literature support and background for electrochemical support removal.469Several groups have shown that secondary materials can be precisely added to the LPBF bed using inkjet technology [17-20]. Using this approach, local additions of sensitizing agents to the support material could be used to create a region that is preferentially etched to free the completed part, reducing or possibly eliminating mechanical machining or the need to co1Tode a shell around the entire part (Figure 2).Figure 2.

Proposed method of support removal by electrolytic etching.MethodsMaterials and FabricationParts were fabricated using a Concept Laser GmbH (Lichtenfels, Gennany) M2 Cusing multi-laser LPBF machine. Mate1ial stock was 316L stainless steel powder (CL 20ES, Concept Laser GmbH) with particles ranging in size from 15 µm to 45 µm. Layer thickness was 50 µm . The sensitizing agent was added to the final 10 support layers of the treated samples. Doping was accomplished by attaching a piezoelectric atomizer or 'mister' integrated with the LPBF coater blade as shown in Figure 3. This approach dopes a strip of the powder for the layers where it is activated. While this does not provide the spatial control of inkjet printing, it simulates the resulting carbon deposition and was more readily integrated into the commercial system.Coater BladeABDESubstrateSubstrateSubstrateSubstratlFigure 3.

Proposed mechanism of precisely depositing the sensitizing agent on the powder bed to cause co1Tosion of the support material. A) A flat substrate is prepared by machining or spreading previous powder layers. B) The coater blade in the LPBF machine spreads a new layer of powder.470C) Deposition device attached to the coater blade deposits the sensitizing agent. D) The laser fuses material as controlled by a 3D model. E) The build plate is lowered to allow the next layer to spread. Note that for the current experiment, a piezoelectric atomizer was used to deposit sensitizing agent in place of inkjet printheads.Etch Rate ComparisonThe relative etch rates of doped and undoped regions were measured using sample two blocks doped with the carbon black sensitizing agent in 3 separate regions as illustrated in Figure 4. One block was heat treated at 700 °C for 10 minutes to sensitize the steel with added carbon, while the other was tested as fabricated.

One block was heat treated at 700 °C for 10 minutes to sensitize the steel with added carbon, while the other was tested as fabricated. The blocks were ground down to 1200f grit using silicon carbide g1inding paper with a manual polisher to get a smooth surface.~ -1. 0 cm ---<(1) 1000 µm■Unmodified(2) 500 µm :::■Doped(3) 100 µm :::lBuild Direct ionFigure 4. Sample block used to test the influence on doped layer thickness on etching rate. The direction of coater movement and dopant spreading is perpendicular to the sample as shown (into the page).They were then etched in a 10% Oxalic acid solution to rapidly attack grain boundaries in the sensitized region, while slowly affecting unmodified SS 316L [21, 22]. Each block was submerged in the acid and attached to an anode with 15V potential for 15 minutes.

Each block was submerged in the acid and attached to an anode with 15V potential for 15 minutes. After etching, the blocks were imaged using a Keyence VHX-7000 Digital Microscope, with a surface profile recorded for each block using the z-stack profilometer feature of the microscope at 400X magnification.Support Removal DemonstrationA series of 3 samples with various support types and a porous "Y" strncture were fomrnd with 2 replicates. Supports use either line or solid type parameters, summarized in Table 1, with the geometry shown in Figure 5. The Line, Contour, and Hybrid supp01is attached to their respective parts by 0.1 mm teeth spaced every 1.5 mm. The Tree supports attached with fifteen 0.3 mm diameter branches spaced every 1.5 mm. The Point supports were 0.2 mm diameter and spaced every 3.0 111111. Complete documentation of advanced support parameters is recorded in Appendix C. The dimensions of each sample were measured, and all were found to be within 0.5%.

Complete documentation of advanced support parameters is recorded in Appendix C. The dimensions of each sample were measured, and all were found to be within 0.5%. The471entire build was heat treated at 700 °C for 10 minutes to sensitize the steel. The samples with Tree type supports began to deform after several layers due to thermal stress, so they were aborted mid-print but the partially built samples were still tested with the other samples.Table 1. Scan Parameters used to fonn different suppo1i types and bulksamples.ParameterBulkTypeNIAPower (W)200Velocity(mm/s) 980Spot Si:,e Scan Spacing (um)(urn) 75130SolidLinePoint, Tree, Hybrid Line, Contour370270135012001305090NIAFor electrolytic support removal, the build plate with parts and supports attached was inverted and etched in an oxalic acid bath with + 15V potential for 200 minutes as shown in Figure 5.

After etching, each sample was removed in pure tension by an Instron 3400 Single Column Series tensile testing machine. Because support dimensions were unknown after etching, only total load to remove the parts was measured for comparison between doped and undoped supports.Figure 5. Diagram of experimental build plate with supports inverted inan electrolyte bath for electrolytic etching.Results and DiscussionSensitization of Doped RegionsEtching created distinct grooves (up to 100 µm deep) in the doped regions of the test blocks, while the unmodified material maintained a flat profile. Heat treatment significantly increased the sensitization of the steel, as proven by grooves 2x deeper etching in the heat-treated sample than the as-fabricated sample (Figure 6b). The same heat treatment was therefore applied to suppo1ied parts in proceeding experiments.

The same heat treatment was therefore applied to suppo1ied parts in proceeding experiments. It was also observed472that there was not a strong relationship between the thickness of the doped region and the etched groove depth, indicating that as few as 5 layers of the suppo1i material needs to be doped for preferential etching to occur (see Figure 6).1.0cm11) 1000 µm::: a. (2) 500 µm ::: (3) 100 µm :::UnmodifiedDoped0 .00 .0\ 1Urihc('tcd r e.:a.t Treated/,,,,---~_,.I 1- -=..---.. .,E :i0.05..c n -0.1CIJ Cl0.15E-o.os 2, ..c n. -0 .1 CIJ a0,15--~ / I t \. i I - r ·. / I II ,' ·,~ \ / / ,, b. deta.il)\1 · . , .• · .~, ..... _I / ·-·0 0 .5 1.0 1.5 2.0 2.5 3.0 3.5 4.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75Posit io n (µm )Position (µm)Figure 6. a) Geometry of expe1imental blocks used for validation. Region 1 is 1.0 mm wide, Region 2 is 0.5 1m11 wide, and Region 3 is 0.1 mm wide.

a) Geometry of expe1imental blocks used for validation. Region 1 is 1.0 mm wide, Region 2 is 0.5 1m11 wide, and Region 3 is 0.1 mm wide. b) Groove depth of sensitized regions of varying widths after 15 minutes of etching. detail) Groove depth profile for 1.0 mm wide doped region with and without heat treatment after 15 minutes of etching.Support StrengthA graphical summary of the force required to remove each support type is shown in Figure 7. Sensitization treatment reduced the force required to remove supports in every case, but the change was most pronounced for the Tree and Contour types where the force required to remove parts from their supports was reduced by over 90%. The sensitized sample with the Tree supports separated while attaching the grip for the tensile tester, so the force required to remove it was negligible. No comparison is provided for the Point supports because both the doped and undoped supports separated due to etching, making any comparison trivial.

No comparison is provided for the Point supports because both the doped and undoped supports separated due to etching, making any comparison trivial. Literature shows that the doped mate1ial has increased tensile strength [23]. Thus even the Line supp01is must have had reduced surface area in the sensitized supports to break with lower overall load than the unsensitized control.4731400~ 1200 z1000Q) > 0 E 800 ~ .., 0 Q) 400 ::: 0600u.. 200 -• - ■0ContourHybridTreeLine■ Sensitized ■ UnmodifiedFigure 7. Force required to remove supports, sorted by type andtreatment.Another observation from tensile testing is that the sensitized supports fractured at a predictable location through the doped layers, while the unmodified supports yielded in a ductile way and broke in unpredictable locations, as shown in Figure 8. These failure modes were expected based on previous research showing fusion of mixed-grade steels [24].

These failure modes were expected based on previous research showing fusion of mixed-grade steels [24]. These results suggest that this material system could also work without etching if the parts can support the force required to remove them by simple tension.LinePointTreeUndopedDopedDopedUndoped DopedHybridContourUndopedDopedUndopedDopedFigure 8. Ductile failure of supports fo1111ed by unmodified material andbrittle fracture of doped supp01is for each support type.GeometryOne problem with using corrosion mechanisms to remove metal samples from their supports is that submerging either the finished parts or the build plate in an etching fluid results in material removal from any474submerged surface. This was a challenge of the cmTent experiment.

This was a challenge of the cmTent experiment. Although SS 3 l 6L is resistant to macro etching in oxalic acid, the overall dimensions of the suppmied parts were reduced unifom1ly by 0.25 111111 due to their full submersion and long exposure to the electric potential. This could be reduced by optimizing the potential to maximize the contrast in etching rate between the doped and undoped regions as well as modifying the support parameters to minimize the thickness of the samples at the etch region.General etching of the samples becomes a more obvious problem upon examination of the porous Y structures as in Figure 9. Both Y structures were reduced to thin delicate members during the etching process, despite being fabricated as robust members. If this procedure were used to avoid deforn1ing delicate features during support removal, the integrity of those delicate features could be compromised by exposure to the etching process.UnetchedEtchedFigure 9.

Comparison between unetched and etched samples, both after heat treatment. The Y structure obviously contains thinner members after etching, which is indicative of the change in geometry due to etching.As a simple illustration, a second experimental build was perfom1ed with similar support types. In the second experiment, rather than inve1iing the build plate to suspend the parts in etching fluid, the build plate was coated with an insulating polymer and fully submerged in the etching fluid so the paiis would not etch. The results, shown in Figure 11 and Figure 11, indicate that the modified etching configuration not only protects the structure of the printed parts, but also enhances etching in the supports, as less etchable surface was exposed to the electrolyte bath and the parts separated without any tension.4751200~ 1000 ~800Q.J > 0 E ~ 0 .., u L. 0 u.. 200600Q.J 400 I0Block■ - ■ - HybridTree■ Sensitized ■ UnmodifiedFigure 10.

0 u.. 200600Q.J 400 I0Block■ - ■ - HybridTree■ Sensitized ■ UnmodifiedFigure 10. Force required to remove supports with the build platesubmerged in the acid bath rather than the paiis.Other potential solutions to the problem of part etching dming supp01i removal are to use an etchant that more aggressively attacks the sensitized material for a shorter amount of time, such as the Al-7 etchant suggested by Nesbitt et al. for macro-etching of ferrous alloys [25] or to use a flexible cathode wire which could be strategically positioned near the supports to be removed.C.a.Figure 11. LPBF parts with sensitized supports released from the buildplate during etching without need for any applied force.476ConclusionsEn-situ powder bed doping was employed to sensitize support structures for LPBF of stainless steel 3 l 6L. Despite using a gentle etchant, the dimensions of the product were altered by the etching process.

Despite using a gentle etchant, the dimensions of the product were altered by the etching process. However, the sensitized supp01is demonstrated a predictable fracture location when pulled in tension, unlike unmodified supports which yield and break unpredictably. Etching also reduced the dimensions of the sensitized supports sufficiently to decrease the fracture load by over 90% for certain support types.

Macromolecular Symposia, 2022. 404(1). Schmithusen, T. and J.H . Schleifenbaum. Wet-chemical support removal for additive manufactured metal parts. in 30th Annual International Solid Freeform Fabrication Symposium - An Additive Manufacturing Conference, SFF 2019, August 12, 2019 - August 14, 2019. 2019. Austin, TX, United states: The University of Texas at Austin. Nisser, M., et al. Sequential Support: 30 Printing Dis.solvable Support Material for Time-Dependent Mechanisms. in TEI '19: Thirteenth International Conference on Tangible, Embedded, and Embodied Interaction. 2019. ACM. Lefky, C.S., et al. Dis.solvable metal supports for printed metal parts. in 27th Annual International Solid Freeform Fabrication Symposium - An Additive Manufacturing Conference, SFF 2016, August 8, 2016 - August13.48114.15.16.17.18.19.20.21.22.23.24.25.10, 2016. 2016. Austin, TX, United states: The University of Texas at Austin.

rmulation Of Lamination Conditions And Interface Studies Using Acrylic Binder System For LOMAnirban Das§ , Niladri Dasgupta#, B. Gurumoorthyk , A. M Umarji§ § Materials Research Centre, Indian.Institute of Science, Bangalore-560012, India k Department of Mechanical Engineering, Indian Institute of Science, Bangalore-560012, India #Ceramic Technological Institute, BHEL,Bangalore-560012, IndiaAbstract:A multicomponent binder-plasticiser system based on acrylates was formulated for making flexible tapes of Alumina (58 volume %), to be used for Laminated Object Manufacturing. Optimum lamination parameters were arrived at by viscosity measurement of the binder-plasticiser system and dilatometry of the compounded tapes . Lamination was done at five different temperatures (75, 85, 90, 95 and 110 (cid:176) C) with solvent (toluene-butanol) spray and under three different pressures ( below 50 psi).

Lamination was done at five different temperatures (75, 85, 90, 95 and 110 (cid:176) C) with solvent (toluene-butanol) spray and under three different pressures ( below 50 psi). The interlaminar strength was measured and interface studies were made using SEM .The shrinkage along the Z direction during lamination was in the range of 1.7-3.5%. Highest interlaminar strength and defect free lamination could be achieved at the intermediate temperature of 90oC, due to sufficient reduction of polymer viscosity and presence of solvent.Introduction:Laminated Object Manufacturing (LOM) is a layered manufacturing technique which is used to create solid 3D shapes of materials by sequential cutting and laminating 2D sheets or tapes of the same material. The principle behind this is essentially same with most of the other Rapid Prototyping techniques i.e.

The principle behind this is essentially same with most of the other Rapid Prototyping techniques i.e. to fabricate components directly from CAD solid models .Ceramic tapes are made and cut in specific CAD defined contours in 2D and ultimately laminated to realise a 3D shape[1].

to fabricate components directly from CAD solid models .Ceramic tapes are made and cut in specific CAD defined contours in 2D and ultimately laminated to realise a 3D shape[1]. The materials issues are to make suitable flexible tapes as feed material, formulation of a thermoplastic organic system for the making of these tapes, formulation of optimum lamination conditions and interface studies, and finally binder burnout and defect free sintering.The LOM technology has been successfully used in making complicated shapes of different ceramics (Alumina, Silicon Carbide, Silicon Nitride etc) and composites [2].While considerable reports dealing with the lamination process of Alumina and other ceramics are available [3] , not much details are available with respect to binder formulation, solids loading, viscosity of the polymer system and their effect on lamination conditions and formation of defect free good interface. LOM was tried to make shapes of alumina.

LOM was tried to make shapes of alumina. We have formulated a specific binder-plasticiser system based on polyacrylates to make alumina tapes.The selection of the organic system was based on the thermal properties of the organic polymers. Our objective was to formulate a suitable lamination condition for LOM based on this binder system. Optimisation of lamination conditions by investigating viscosity of the organic system as a function of temperature, dilatometric studies for shape integrity during binder burnout, interlaminar strength and interfacial defect elimination are reported here.9Experimental:A tape casting slurry of Alumina ( ACC-India A-16 grade, average particle size:0.6 microns, specific surface area:8-10 sq.m/gm) was made using an acrylate binder and polythelene glycol and benzyl butyl phthalate as plasticiser in toluene- butanol solvent (in azeotropic ratio) medium using gleceryl trioleate as dispersant.

The alumina powder dried at 100-110(cid:176) C for 30 minutes was mixed with the dispersant which was disssolved in the solvent system along with 0.25 %by weight talc used as grain growth inhibitor. It was then milled in a pulverisette ( speed 144 rpm)with high purity alumina as grinding medium. for 3 hours. Then the binder and the plasticisers were added to form a viscous slip. Once again it was subjected to milling in the pulverisette for 3 hours. The slip was vacuum degassed and sieved through a nylon mesh to remove any large particles or undissolved binder, if present. The viscosity of the slip was measured to be 1790 cps in a Brookfield viscometer (DV-II type, spindle no. 25) at 60 rpm. It was then cast in a batch type double-blade caster (casting rate is 10 cm/min) onto a silicone coated polyester film . Then the tape was subjected to drying for 10 hours, the tapes of thickness 0.55mm were peeled off the film and cut to desired shapes for subsequent measurements.

Then the tape was subjected to drying for 10 hours, the tapes of thickness 0.55mm were peeled off the film and cut to desired shapes for subsequent measurements. The solid loading was 58% by volume with respect to the total organic binder content after solvent removal.Table 1: Slip formulation used for Alumina tape castingConstituents Ceramic Powder Solvent Binder Plasticiser I Plasticiser II DispersantMaterials Alumina Butanol-Toluene mix Polymethyl Methacrylate Polyethelene Glycol Pthalete Gleceryl TrioleateVolume % 29.52 49.14 9.08 3.50 5.26 3.50The viscosity of the binder-plasticiser-dispersant system was measured in rotational viscometer (Brookfield Viscometer DV-II type, spindle no. 32 ) at a temperature range of 35- 100(cid:176) C at 8 different temperatures and at different rpm’s at each temperature. Lamination was done using solvent spray [4] and by the application of pressure and temperature. Tapes were cut into rectangular pieces of 1cm · 1 cm.

Lamination was done using solvent spray [4] and by the application of pressure and temperature. Tapes were cut into rectangular pieces of 1cm · 1 cm . Lamination was done at 5 different temperatures (75, 85 ,90 ,95, 110(cid:176) C) and under three different pressures (1.75 , 2.2, 3.0 kg/sq.cm). Two tapes of the said dimension were joined together after solvent spraying and then placed in an apparatus fabricated in-house for lamination. The tapes were kept in the particular pressure-temperature condition for 10 minutes. The change in the thickness of the laminates parallel to the pressure axes ( the z direction) after lamination were measuredThe interlaminar strength was measured using an arrangement as shown in fig.1.Two laminated tapes, after lamination using different pressure-temperature conditions were used. Each face of the laminate were stuck to two identical rectangular brass plate assemblies using a two component epoxy resin.

Each face of the laminate were stuck to two identical rectangular brass plate assemblies using a two component epoxy resin. The brass plate assembly (fig.1) was hung and the load applied. The10particular load at which delamination occurs was noted. The laminate interfaces were studied using SEM (S360 Cambridge Scanning Electron Microscope) after gold plating.Dilatometric measurements were made in a laminated sample (height 4.85 mm). Thesample was heated to 800(cid:176) C at 2(cid:176) C/min in the dilatometer and then cooled to room temperature.Brass assemblyLaminatesLoadFig 1: Experimental arrangement for interlaminar strength measurementResults and Discussions:Flexible tapes of uniform thickness (0.55 mm ) were obtained after casting followed by drying. The tapes had a good surface finish and were easy to handle.

The tapes had a good surface finish and were easy to handle. The loss on ignition showed the total organic content was 16.48%, which corresponds to 41.96% by volume of the starting formulation without the solvent.Fig 2: Variation viscosity of the organic binder-plasticiser system with temperature at different shear rates11The viscosity data of the organic components show that viscosity reduces sufficiently after 60C (shown in fig.2). Viscosity comes down most significantly in the range 75-100(cid:176) C, and after 100(cid:176) C viscosity came under the instrument detectable range. Also the viscosity shows increase with decreasing shear, which implies pseudoplastic behaviour of the organic system.

Also the viscosity shows increase with decreasing shear, which implies pseudoplastic behaviour of the organic system. From the viscosity data lamination temperature range was fixed to be between 75-110(cid:176) C.Attempt was made to find out the viscosity of the compounded ceramic-polymer mix using a capillary rheometer (CFT-500C type, Shimadzu Corporation) with capillary diameter 1mm, but no signifiacant sagging and softening was observed till 200(cid:176) C and a load of 500 KN.The lamination was to be done on green ceramic tapes, so the pressure to be applied depended on the allowable dimensional change of the laminate after laminating and retention of shape of the green tape. After laminating at different lamination pressure-temperature conditions it was found that the dimensional change along the pressure axis (Z axis) was between 1.7-2.6% without deshaping of the laminates ( Table 2).

The strength increased with temperature till 90(cid:176) C and then decreased . With increasing the pressure during lamination the strength data shows increase for all the temperature conditions. Strength was found to be highest at a pressure of 3.0 kg/sq.cm. The SEM figures (at three different temperaures at a load of 3.0 kg/sq.cm) also showed the best interface was observed in the pressure-temperature condition of 90(cid:176) C and 3.0 kg/sq.cm (Fig. 4 b). Both at higher and lower temperatures showed clear clear delamination (Fig 4a, 4c). At lower temperatures the organics did not soften enough to give good lamination even at higher pressures.

4 b). Both at higher and lower temperatures showed clear clear delamination (Fig 4a, 4c). At lower temperatures the organics did not soften enough to give good lamination even at higher pressures. At temperatures above 90(cid:176) C due to evaporation of solvent ( azeotropic b.p is 105(cid:176) C)12Fig 3:Interlaminar strength Vs lamination temperature at different lamination pressures(a)(b) 500m m (c)Fig: 4 :SEM micrograph of interfaces obtained at 3 kg/sq.cm and at different temperatures (a):110(cid:176) C (b): 90(cid:176) C (c):75(cid:176) C13lamination was poor.

Though both the strength data and SEM studies showed with increasing pressure during lamination ensures better interface, but due to the factors like deshaping of the tapes and significant dimensional changes during lamination the pressure was limited to the value of 3.0 kg/sq.cm.The result of the dilatometric measurements of a laminated (at 90(cid:176) C and 3.0 kg/sq.cm) sample of height 4.85 mm is shown in Fig.5. There is no significant dimensional change or shape loss in the sample till 450(cid:176) C. On further heating the contraction due to the binder softening and decomposition is observed. During cooling the sample showed the usual contraction of a porous ceramic sample ( a = 4.2 · 10 –6/K).

On further heating the contraction due to the binder softening and decomposition is observed. During cooling the sample showed the usual contraction of a porous ceramic sample ( a = 4.2 · 10 –6/K). This shows that our binder formulation is suitable for making complex ceramic shapes without need for creating support structures.Fig 5: Thermal strain of the laminated alumina body as functionof temperature during heating and coolingThe binder burnout studies, retention of net shape and studies on laser cuttability of green ceramic tapes are currently under progress.Summary and Conclusion:Flexible alumina tapes were produced an acrylic binder and a non-aqueous solvent system. Based on the viscosity of the organic system at different temperatures lamination temperature was chosen. Lamination was done using solvent spray and under applied load. Optimum lamination conditions were arrived at by measuring the interlaminar strength and studying the interface using SEM.

Lamination was done using solvent spray and under applied load. Optimum lamination conditions were arrived at by measuring the interlaminar strength and studying the interface using SEM. Defect free strong interface was achieved at a temperature of 90(cid:176) C and under a load of 3.0 kg/sq.cm. It was seen that with increase of load interface was better but dimensional change in the Z direction was more with increasing load. At temperatures less than 90(cid:176) C the organic system was not soft enough to give good lamination and at temperatures higher the sprayed solvent evaporates as its boiling point is 105(cid:176) C.14References:1.2.3.4.5.C. Griffin, J. Daufenbach, S. McMillin, “Desktop Manufacturing:LOM vs Pressing” American Ceramic Society Bulletin vol 73(8), 109-113 (1994) Z.E Liu, P. Wie, B.Kerman, A.H. Heuer, J.D.Cawley, “Metal and Ceramic components using CAM-LEM Technology” Solid Freeform Symposium 1996, 377-384 J.D Cawley, Z.

ACTAL GROWTH MODELING OF ELECTROCHEMICAL DEPOSITION IN SOLID FREEFORM FABRICATIONJack G. Zhou, Zongyan He and Jian Guo Department of Mechanical Engineering and Mechanics Drexel University, Philadelphia, PA 19104ABSTRACTA new rapid tooling technique ElectroChemical Liquid Deposition Based Solid Freeform Fabrication (ECLD-SFF) was introduced in this paper. In the ECLD-SFF a substrate made of or coated with conductive materials is connected to a DC power supply, and the substrate is put into a plating bath. A very thin metal pin is connected to the DC power as a positive electrode. Between the substrate and the tip of the pin there is a thin layer of metal powder. Under the effects of electric field, metal ions from electrode moving to chemical liquid will deposit onto the powder particle and growing so that the metal particles can be bound by the deposited materials to form freeform solid.

By controlling the pin's movement and electrified time, a desired 3-D shape will be built through layer by layer scanning. distinguishes itself from other SFF techniques with advantages of products: high build rate, high accuracy, high density, low shrinkage and controllable microstructures. It has been found that the electrochemical deposition among metal particles during ECLD-SFF is a fractal growth process. The fractal dimension and the width of the deposited metal band are all related to electric field density, composition of Several models on the fractal growth between electroplating liquid and processing time. electrodes or metal particles were developed in order to explain these fractal growth phenomena and obtain desired process parameters and conditions for the ECLD-SFF process.INTRODUCTIONThere have been many.efforts to use gas-phase laser-induced reaction to make parts.

The Selective Laser Sintering system has been usedto form alumina parts by oxidation of aluminum powder in air, and to form silicon nitride by laser sintering .of silicon in nitrogen or ammonia (Birmingham, 1995). Other efforts are focused on converting laser-induced Chemical Vapor Deposition (CVD) from a film-forming to shape-making technique (Maxwell, 1995, Lehmann, 1994). A technique called Selective Area Laser Deposition (SALD) has been proposed by Jakubenas et aI., (1997). Their proposed process, depositing material from one or more organometallic gases on a. substrate selectively heated by a scanning laser beam, offers an opportunity of forming shapes under lower temperature with a wide range of materials. The primary limitation of gas-phase reaction methods is their low deposition rates.

The primary limitation of gas-phase reaction methods is their low deposition rates. According to the listed deposition rate of organometallics, 0.1-1.0 ,..un/min, (Sudarshan, 1989),it may take several weeks even months to make a cubic inch solid metal part by using conventional CVD. Solids can also be free-formed by electrochemical methods. There are two SFF techniques related to electrochemical reactions: tooling with Nickel Ceramic Composite (NCC) and Expresstool (Ashley, 1998). The NCe tooling method uses plastic RP models as a master pattern, it is first coated with a conductive silver-based material, then placed in an electroforming bath of nickel sulfamate where a thin nickel layer· is plated over it. The typical nickel plating thickness varies from 0.04 to 0.2 inch (Ashley, 1997). After electroforming, Chemical Bonded Ceramic (CBC) is cast to support the nickel shell. Once the CBC is cured, the ejector pins are drilled and installed.

After electroforming, Chemical Bonded Ceramic (CBC) is cast to support the nickel shell. Once the CBC is cured, the ejector pins are drilled and installed. This technique is not a direct free-forming method, and the filled-in ceramic material usually has229large shrinkage that will cause shear stress on the interface and tooling distortion. In Expresstool technique, a mandrel is first made with a CNC machine, then the mandrel is put in a bath of nickel sulfamate for electroforming. The formed nickel shell with a thickness of 1 to 2 mm needs to be backed withaluminum...filled epoxy. The aluminum fraction helps heat conduction. Two main advantages of this process are the ability to produce large parts and the high accuracy of the products due to a very small shrinkage. A key disadvantage is that deep holes do not electroform well; this is also the shortcoming of conventional electroforming techniques caused by the heterogeneity of electric current density.

A key disadvantage is that deep holes do not electroform well; this is also the shortcoming of conventional electroforming techniques caused by the heterogeneity of electric current density. This technique is also not a direct freeform fabrication. A new SFF technique has been developed to overcome the main limitations of other gas-phase chemical reaction or electrochemical SFF techniques. It will be introduced briefly in this paper.fractal geometry (Mandeibrot, 1967, 1982) has blossomed tremendously in the past few years and has helped reconnect pure mathematics research with both the natural sciences and engineering. Within the last decade fractal geometry and its concepts have become centraltQolsin most of the natural sciences, including physics, chemistry, biology, geology, and meteorology. The application of fractal geometry in engineering,however, is still in its infancy.

The application of fractal geometry in engineering,however, is still in its infancy. Although in its infancy in engineering practice,Jractal geometry has already been introduced into many disciplines· of engineering research and it has great potential in developing new technologies and solving real engineering problems. The engineering field that has benefited the most from fractal research is materials science. The microstructure of such materials as metals, composites, concrete, rocks, graphite, and the majority of synthetic materials can be characterized by fractal geometry. Particle distributions and aggregations also possess; fractal properties. The •. fractal characterization of materials has greatly· helped researchers in understanding the microstructures. and operative mechanisms of materials,· and in implementing the manufacture. of new materials. In this paper, the authors will use fractal geometry as a to study deposits ·and.

and operative mechanisms of materials,· and in implementing the manufacture. of new materials. In this paper, the authors will use fractal geometry as a to study deposits ·and . particle growing mechanism in the modeling and analysis tool electrochemical liquid deposition process for solid freeform fabrication. If the fractal growth process is realized and understood thoroughly, then it can be used to predict and control various electrochemical liquid deposition processes used in coating, plating, solid freeform fabrication, and electronic circuit and device making (Bradley et aI., 1997).On the other hand,A NEW SOLID FREEFORM FABRICATION TECHNIQUEA new SFF technique, called ElectroChemical Liquid. Deposition based SFF (ECLD SFF), has been developed by the authors to overcome thernainlimitations of other gas-phase chemical reaction or electrochemical SFF techniques.

Deposition based SFF (ECLD SFF), has been developed by the authors to overcome thernainlimitations of other gas-phase chemical reaction or electrochemical SFF techniques. In the ECLD-SFF (see Fig.l), the substrate is made of or coated with conductive materials (metals or graphite), and is connected to a DC power supply as a negative electrode (cathode). Then the substrate is put into a plating bath that is filled with electroplating liquid. A very thin pin that is made of deposition metal is connected to the DC power as a positive electrode (anode). Between the substrate and the tip of the pin there is a thin layer of metal powder/particles. Between cathode and anode there is an electric field named Z direction field. Two assistant electric fields are arranged perpendicular to each other to form an X-Y surface electric field. A magnetic field is applied in the Z direction of the substrate to form a tight connection of ferrite material powders.

A magnetic field is applied in the Z direction of the substrate to form a tight connection of ferrite material powders. Under the effects of the electric and magnetic fields, metal ions from electrode moving to chemical liquid will deposit onto the powder particle and growing, and the metal particles will be bound by the deposited materials toform freeform solid. By controlling the pin's movement and electrified time, a desired 3-D shape can pe built through layer by layer scanning. The formed product will be further treated, such as sintering and infiltration. Fig. 2 is a conceptual diagram of the ECLD-SFF system.

2. The conceptual diagram of the ECLD-SFF systemThe new ECLD-SFF distinguishes itself from other SFF techniques with the following advantages: (1) High build rate. In ECLD-SFF, the deposited material serves as the binder among the particles of metal powders, and the prodqct is fonned by many layers of thFPowders not only the deposited materials, thus a high producing rate can be expected. Additionally, the rate can be increased greatly by changing the shape ofthe.anode. For example, for a part with large flat surface we can choose a similar shape metal plate as an anode. (2) High accuracy. In ECLD-SFF, a specially designed insulation tube is used to constrain the electric field of the pin231anode, which will provide a concentrated electric field between the substrate and the pin so that the deposition occur only in a very small area. Since both the tube and pin wire can be made very thin,. the scanning accuracy will be very high. (3) High density and low shrinkage..

Since both the tube and pin wire can be made very thin,. the scanning accuracy will be very high. (3) High density and low shrinkage.. In the assistant electric fields are used to increase the amount of the deposited ECDB-SFF, materials among the powders, and the assistant magnetic field is used to increase the accumulation density of the ferromagnetic powders. Also, the layer thickness· of the powder can affect the density. The thinner the layer of the powder, the higher the density should be. The (4) Controllable lead low shrinkage in later sintering and infiltration. high density will In ECLD-SFF, the microstructures of the formed part can be controlled by microstructure. changing the composition and the layer thickness of the powder.

high density will In ECLD-SFF, the microstructures of the formed part can be controlled by microstructure. changing the composition and the layer thickness of the powder. For example, for the working surfaces of a mold, which needs a higher hardness, we can reduce the layer thickness of the powder in order to increase the density or we can use a mixture of ceramic and metal powders.FRACTAL PHENOMENA IN ECLD-SFFExisting researches on electrochemical deposition and our initial experiments on ECDB SFF revealed the following interesting phenomena and also challenges. (1) Researches and experiments have shown that metal particles in a suitable electroplating medium can be connected by directed electrochemical growth (Bradley, et aI., 1997, 1998). Our experiments further show that if some metal powders are laid on the cathode plate (substrate) and sunk into the liquid medium (Fig. 1), the powder particles can be bound under the action of electric fields In Fig.

1), the powder particles can be bound under the action of electric fields In Fig. 3a, an SEM photo shows a ropelike deposit and form a solid with a certain strength. connecting to a particle at the initial growth stage. Our research result, in Fig. 3b, shows the formed powder layer bound by the deposits. (2) Researches and experiments have shown that the electrochemical deposition is a fractal growth process even the distance between two particles is very short, and the fractal growth always occurs in the direction of the electric field (Brady and Ball, 1984). The fractal dimensions of the growing branches and the width of the deposited metal band all decrease with the increasing electric field. Fig. 3c shows several stages of the fractal growth, from which one can see when some new fractal branches are forming, some formed fractal branches may disappear at the same time. After a longer time only a few main branches can survive.

After a longer time only a few main branches can survive. (3) Researches and experiments have shown that by using assistant electric fields in surrounding directions (Fig. 1), a web-like deposits among particles can be formed (Fig. 3d). (4) Our initial research has shown that the surface topography of the bound powder particles has fractal structures, and the fractal dimension depends mainly on the composition of the medium, the processing time and the strength of the electric and assistant magnetic fields (Fig. 3e). An assistant magnetic field can increase the particle density and affect the dimensions of the (5) Researches and experiments have shown that during the electrochemical fractal surfaces. process the voltage and current are unstable. There is a critical point of processing time for the deposition among each layer (see Fig. 4). At this critical point the constant voltage will drop sharply to a very low value.

First, we consider the simplest case i.e., the fractal growth between two electrodes, then consider the model between233two powder particles in single electric field. For the growth among thepowd~1"scontaining a large amount of particles, we will try tou.se the re-normalization group method.The goal of this model is to find the density distribution ofthedepositinia small area between the pin f}nodeandtheplate cathode. It is expected that the density distribution is related to the microstructures and accuracy ofthe product. The pin anode with a diameter d is above the center of a metal disc and the cathode/disc has a diameter D. We assume that d is much smaller than D, and the end of the pin is so.close to the disc surface that the deposition can be assumed as a 2-D process occurring on the disc. After a metal atom in the pin anode losses its electrons and become an ion, it will move into the plating liquid.

After a metal atom in the pin anode losses its electrons and become an ion, it will move into the plating liquid. Once the ion contacts thegateode surface it will get electrons and deposit on the disc. The second, ion wal.ks randomlYii~ntilit visits a site adjacent to the first one and the walking particle becomes part of the cluster. >The next particle joins the cluster at a random distant point, and so forth. With the increase of deposited atoms the average walking distance of an ion before contacting the cathode will decease. The exposed ends of the cluster tend to grow more rapidly than other perimeter sites beca~seitbese sites are "shadowed" by the deposits. The above mentioned growth process can besirn.l.llated with the Diffusion-Limited Aggregation (DLA) model. This model was first developed by Witten & Sander (1981), inwhichthe so-called "shadow" effect is expressed as "diffusion-limited' In a 2 D DLA model a particles, i.e.

This model was first developed by Witten & Sander (1981), inwhichthe so-called "shadow" effect is expressed as "diffusion-limited' In a 2 D DLA model a particles, i.e. ion is released from a random positionon the bounclarY of a square lattice at regular intervals. The released particles will walk randomly in theilatticeuntil join the aggregate. If assuming the distance y separating two sites is much less than the size D of the aggregate, we can obtain information about the deposit-particle distribution from the following density correlation function:C(y) = N- 1LP(Y')P(Y'+ y),(1)y'where y' denotes the distance variable, N is the total number of deposit-particles, the density p(y) is defined to be 1 for the occupied site and 0 for the others. Fig. 5a shows a 3600-particle aggregate on the lattice. Fig. 5b shows the simulation results, in which the solid line is a least squares fit over the range y= 3 to y = 27.

Fig. 5a shows a 3600-particle aggregate on the lattice. Fig. 5b shows the simulation results, in which the solid line is a least squares fit over the range y= 3 to y = 27. The error bars representthe spreads of values among the six salnples of aggregate. The arrow marks the average radius of the gyration. The simulation showsC(y) =0.43y-0.343±O.OO4 •(2)From this we can get the following results. (1) The density distribution depends only on the distance separating two sites providing the distance is much less than the size of the aggregate. (2) The density correlation within the model aggregation falls off with distance obeying a fractal power law. (3) The maximum of density is only 0.43 (let y=1 in Eq. (2)). The lower density of the deposit will cause large shrinkage, distortion and low strength of product. This is why we do not use direct electrochemical deposition between two electrodes.

(2)). The lower density of the deposit will cause large shrinkage, distortion and low strength of product. This is why we do not use direct electrochemical deposition between two electrodes. (4) Considering that the real deposition occurs on a smaller circular area not a large square area and the ions is not released from the boundary of the deposition area, the density of the center region close to the pin anode must larger than that of others. This kind ofdistribution is desirous to the high accuracy of products. This 2-D model can be analogized to 3-D and was verified by experiments (Brady, 1984). However, DLA model is not so suitable to describe the fractalgrowth in our ECLD-SFF study. First, according to DLA model, we can not get any formula used to calculate the growth rate. Second, our experiments showed that the fractal growth occurs between the polarized powder particles not between the electric poles.

Second, our experiments showed that the fractal growth occurs between the polarized powder particles not between the electric poles. Thus the space between particles, the sizeiand shape of the particles all affect the growth. Thirdly, according to the DLA model, the growth process can continue forever without any critical points of cease-growing like illustrated in Fig. 4. Therefore, othertwo models were developed.2. Modeling on the Fractal Growth between Two Metal ParticlesThe main goal of this model is to. obtain a formula on the growth rate of deposits between powder particles. If the boundary and initial conditions can be got, according to the formula we Production rate is very can calculate the growth rate, important not only to ECLD-SFF but also to other rapid tooling techniques. The fractal growth between two powder particles (not atoms) is shows in Fig. 6. After an electric field is applied to a pair of particles, both of them will be polarized (Fig. 6a).

The fractal growth between two powder particles (not atoms) is shows in Fig. 6. After an electric field is applied to a pair of particles, both of them will be polarized (Fig. 6a). Initially the particle close to the anode will liberate metal ions while the particle close to cathode will reduce solvent (Fig. 6b). The concentration of the ions near the former one is much higher than that near the latter one. The ionswiU.move towards the latter one by diffusion, and when the ion concentration around the latter particle reaches a critical value the electrodeposition will occur and the fractal branch begin to grow to facing the other particles (Fig. 6c). It is first assumed that the single electric field between two spherical particles has a cylinder shape (Fig. 6d), the ion concentrated along the X-axial direction has a continuous distribution. If the.

6d), the ion concentrated along the X-axial direction has a continuous distribution. If the. distance between the two particles is long enough, the effect of the shape of the particle can be neglected, then the concentration distribution along the radial r-direction will be uniform. Thus we can use U(x, t) denotes the In a concentration at point x at time 1. IfR denotes the radius of the particle, we have -R< r <R. very small area O(x, r) around the center point C(x, r), the ion number will be U(x, t)o(x, r). If assuming· that ions all obey Brownian motion law and are independent with each other, then it can be deduced that the ion concentration at time t + h has the following normal distribution:then estimate the production rate.U(x', t + h) =(21thr 1l2 exp[-(x - x,)2/(2h)]U(x, t)dx By differentiating x' and h, we obtain au/at =0.5 V2U(3)(4)where V2 denotes the second order partial differentiation.

The boundary conditions of this differential equation are assumed as follows: On the surface of the particle near the·cathode the ion concentration is a constant during the process, i.e., U(x = 0) = Uo. Using Ft denotes the depositing bound attime t, on Ft all ions will lose their electric charges, so We have U(FD = O. The growing rate of Ft in its orthogonal direction will beVn =Kn VU(5)where V denotes the first order partial differentiation and K is a coefficient of per unit time growing rate.This model also can not describe the critical phenomenon shown in Fig. 4, because it is related to only a few particles. As we know most critical phenomena occur in the systems consisting of a large number of individual units.235o~oa)O=~~ Q:~;.~~rt-+E1U (x,t)xc)metal atom metal iond)Fig. 6. Fractal growth of deposits between two parcicles3.

6. Fractal growth of deposits between two parcicles3. Modeling on the Fractal Growth among a Large Amount of ParticlesIt is impossible to treat fractal growth related to many particles by conventional analytic methods. Even if we can deduce a very complex model to describe the growth among ten particles, we are still not sure if it can be used to describe the case of more than ten particle, and whether the results can be used to explain the complex growth phenomena. So we wilItry to use the re-normalization group method which has been used to treat various phase transformation or critical problems (Wilson and Kogut, 1974). Considering a 2-D array which contains a large amount of particles, if four neighbor particles can be connected by deposited material just like that shown in Fig. 3, we call the unit containing four connected particles breakover cell. Under the action of the electric fields, some units in the array may become breakover cells, but others may not.

3, we call the unit containing four connected particles breakover cell. Under the action of the electric fields, some units in the array may become breakover cells, but others may not. With the time increasing, the number of breakover cell will be more and more. But when the number reaches a critical value the whole array will become breakover by connecting a lot of neighbor breackover cells and forming some breakover cell clusters. Once the array becomes breakover, the distribution of electric fields on the array will be changed greatly. Further depositing new material among the remained individual particles will be very difficult, so the process will stop. We think it may be the main reason why the voltage will drop down very quickly after a period in the ECLD-SFF process in Fig. 4. The goal of this model is to find the conditions that the cluster can form and the density of the breackover.

4. The goal of this model is to find the conditions that the cluster can form and the density of the breackover. If the formed clusters have the fractal structure, we also want to know their fractal dimension.In the 2-D array model, assuming the breackover probability of a four-particle unit is Po, critical point probability Pc. It has been known when 0 < po < pc, the cluster forming probability P is very small. However, when Pc < po <1, P will be very close to 1. Although by using Monte Carlo method one can calculate the value of pc. when there are a lot of particles, the calculation will be very difficult and time-consuming. Here we are introducing the re-normalization group method. As shown in Fig. 7a, four adjacent particles form into an initial unit, and it has the breackover probability Po. Four adjacent initial units form into a first degree unit with breackover probability PI, and four first degree units form a second degree unit with probability pz, and so on.

Four adjacent initial units form into a first degree unit with breackover probability PI, and four first degree units form a second degree unit with probability pz, and so on. Based on the principle of re-normalization group, pz can be calculated from PI, P3 can be calculated from pz, P4 can be calculated from P3, and so on. Fig. 7b-7f show all probable cases of four first degree units. The probability that all four initial units are not breackover is (1- PO)4 (Fig.7b). The probability that only one initial unit is breackover will be Po(l-Po)3 (Fig. 7c). The probability that only two initial units are breackover should be poz(l-Po)z (Fig. 7d). The236probability that only one initial unit is not breackover is Po\l-po) (Fig. 7e). The probability that all initial units are breackover is P04.

7d). The236probability that only one initial unit is not breackover is Po\l-po) (Fig. 7e). The probability that all initial units are breackover is P04. Therefore, the probability that a first-degree unit is breackover will bePI = 2 po 2(1 - Po )2 + 4 po 3(1 - po ) + Po 4= 2 Po 2- po 4 Then considering the second degree unit with the same method, we get P2 =2 PI 2- PI4. A(6)general equation isPn+1 = 2 pn 2- pn(7)Figure 8 shows the iteration relationship between pn+I and pn. This equation can be rewritten asf(x) = 2x2 - x4(8)p1p2/ ±f p3p4Eal!l EI ~ IiIIl ~ III ~ 8!l II I!III 8Il lIB ~ ~ IIIa}b)c)d)e)f)P.Fig. 7. The breakover probabilities of the first degree unit, Fig. 8. Iteration between Pn and pn+lLet f(x) = x (0< x <1) and we get three solutions (fix points): x l= 0, x, = 0.618 and x, = 1. Their corresponding values of df(x)/dx will be 1..1= 0, 1..,= 1.582 and 1..,= 0.

Their corresponding values of df(x)/dx will be 1..1= 0, 1..,= 1.582 and 1..,= 0. Therefore, both Xl and X, are stable, but x, is unstable because A, >1. So the critical probability p, will be 0.618. It is an approximate value. In order to improve the accuracy we can chose 3 x 3 unit instead of 2 x2unit and obtain,Pn+l = 3 pn 3+ 3 pn 4- 2 Pn 5 - 15pn 6 + 18 Pn 7 - 7pn 8 + Pn 9 According to this equation, we calculate pc =0.609 that is closer to the value obtained from Monte Carlo method. From the obtained critical value, we can deduce the critical particle number ne of the cluster,(9)nnO.5d fnO.5d f(10)where n is the total number of the particles in a 2-D array, and dr denotes thefractaYdimensionof the cluster. If n is big enough we can get dr ~ 1.896. Then the ratio betweellne andnca~be calculated. For example, when n =10,000, the ratio is about 55%.•·It means whenthedypositing process stops, over a half of the metal particles can be connected.

Then the ratio betweellne andnca~be calculated. For example, when n =10,000, the ratio is about 55%.•·It means whenthedypositing process stops, over a half of the metal particles can be connected. However, the real ECLD-SFF process should be considered in 3-D models. In the simplest case, the· initial unit will include 2 x 2 x 2 particles. But there are three ideal stacking modes of the metal particles, i.e., Body237Centered Cubic(I3CC), Face Centered Cllbic(FCC) ... and Rexagonal Close-packed (RCP). For differentstacking. mode the •critical probability wilL be very different. In addition, we must consider the effects of voids in real metal powders. The 3-D model of fractal growth is very challengeand.•promising.CONCLUSIONSIt has/been found that the electrqchemicaL deposition among metal particles during ECLD-SFF is a fractal growth process.

The 3-D model of fractal growth is very challengeand.•promising.CONCLUSIONSIt has/been found that the electrqchemicaL deposition among metal particles during ECLD-SFF is a fractal growth process. The fractal dimension and the width of the deposited metal band are all related to electric Jield density, composition of electroplating liquid and processingtime. Based on the DLAmodel of the fractal growth between two electrodes we know that the.density distribution •• depends only on the distance separating two sites, the density correlation within tlIemodel aggregationfalls off with distance obeying a fractal power law, and the maximum of density is only 0.43. According to a 2..D re-normalization group model, the critical phenomenon on the fractal growth will occur when the whole powder layer becomes breakover.ln the critical case the formed powder cluster's fractal dimension is 1.896, the ratio of the particles connected by the deposits is about 55%.

G. and Kogut, J., "The renormalization group and theexpansion",Phys. Rev. C12, 1974, pp75 Witten, T. A. Jr. and Sander L. M., "Diffusion-Limited Aggregation, a Kinetic Critical Phenomenon",Phys. Rev. Letter, Vol. 47, 1981, pp1400-1403..

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFracture Mechanism Analysis of Schoen Gyroid Cellular Structures Manufactured by Selective Laser MeltingLei Yang, Chunze Yan*, Yusheng Shi*State Key Laboratory of Materials Processing and Die &Mould Technology, School of Materials Science and Engineering, Huazhong University of Science and Technology, Wuhan 430074, ChinaAbstractTi-6Al-4V triply periodic minimal surface (TPMS) structures with biomorphics scaffold designs are expected to be the most promising candidates for many biological applications such as bone implants. Fracture is the main failure mode of Ti-6Al-4V cellular structures at room temperature. However, there is currently less investigation on general analysis about the fracture mechanism of Ti-6Al-4V TPMS cellular structures.

However, there is currently less investigation on general analysis about the fracture mechanism of Ti-6Al-4V TPMS cellular structures. In this work, a typical TPMS structure, Schoen Gyroid, was designed and porous Ti6Al4V Schoen Gyroid specimens were manufactured using Selective laser melting (SLM). Finite element analysis (FEA) method was employed to calculate the stress distribution under compression. The FEA results are used to predict the fracture positions, fracture zones as well as fracture mode. The uniaxial compression experiments were conducted and compared with the FEA results. The experimental and simulation results show high consistency.IntroductionCellular structures have attracted many researchers due to their considerably high performances such as superior energy absorption characteristics, light weight, excellent thermal and acoustic insulation properties and high strength–weight ratio.

A clear description of the mechanical properties is necessary before the specified cellular structure are used in biomedical implants [1] and energy absorption or impact mitigation [2] with their tailored porosity, density, strength, and ductility.Research efforts made in this aspect have been mostly focused on elastic module, yield strength, ultimate tensile strength, which are the main performance parameters of cellular structure materials. Gibson and Ashby are the precursors in this field [3]. They proposed the Gibson-Ashby equations fitted from the experimental data, which show a great potential of predicting these parameters. Maskery [4, 5] investigated the uniaxial compression responses of the functionally graded porous structures of Al-Si10-Mg and Nylon 12.However, Fracture is the main failure mode of the cellular structures fabricated by SLM.

Fractures occurred during the tensile and compression process have a dramatic impact on the mechanical properties of cellular structure. Xiao[6] investigated the2319compression behaviours of Ti-6Al-4V lattice structures at ambient temperature, 200℃, and 400℃, and 600℃ respectively, and found that there are obvious shear bands observed along the 45° plane. The fractures occurred along the shear bands result in a sharp drop of the stress-strain curve and greatly reduce the energy absorption capacity.The structure Optimisation by strengthening the weakness will obviously improve the energy absorption capacity of brittle materials. Hence it is necessary to inquire about the fracture mechanism of these structures, and some research has been done. For example, mechanical behaviours of Al-Si10-Mg lattice structures manufactured by selective laser melting were examined by Maskery et al.

For example, mechanical behaviours of Al-Si10-Mg lattice structures manufactured by selective laser melting were examined by Maskery et al. [4] and fractures were observed almost exclusively at the lattice nodes which are associated with both tensile and shear loading. Mazur et al. [7] investigated the deformation and failure behavior of 6 lattice structures with different cell topologies manufactured by SLM through theoretical prediction and experimental validation, and found that fracture predominately occurs in joints.However, the existing research on fracture mechanism and position are all focused on lattice cellular structures that are derived from Boolean intersections of geometric primitives and have many straight edges and sharp turns. These straight edges and sharp turns will result in highly stress concentration. M.

These straight edges and sharp turns will result in highly stress concentration. M. Smith et al.[8] used finite element models to analyse the stress distribution within the bcc unit cell at increasing levels of crush and found that the formation of plastic hinges in the struts are close to the nodal regions due to the stress concentration. Besides, the horizontal struts in lattice cellular structure will decrease the manufacturability in the additive manufacturing processes[7].Schoen Gyroid is a triply periodic minimal surface structure with smooth infinite surfaces and uniform curvature radius [9]. Without straight edges and sharp turns, Schoen Gyroid cellular structures are expected to show high manufacturability in the selective laser melting (SLM) additive manufacturing processes and excellent mechanical properties [10].

However, there is currently less investigation on general analysis about the fracture mechanism of Ti-6Al-4V Schoen Gyroid cellular structures.In this paper, Ti–6Al–4V Gyroid TPMS cellular structures with an interconnected high porosity of 85% and single unit sizes of 4.5mm were manufactured by SLM. The mechanical properties under uniaxial compression were evaluated by both experimental tests and FEA method. FEA results were used to predict the fracture positions, fracture zones as well as fracture mode.2 Method2.1 Experiment methodThe Gyriod unit cell is mathematically designed by a computational method which is reported in our previous work (Hao et al., 2011). The cellular samples with volume2320fraction of 15% and size of 25mm×25mm×12.5mm were fabricated by 3T RPD Ltd. UK using DMLS EOSINT-M270 machine supplied by EOS GmbH, Munich, Germany.Uniaxial compression tests were carried out using an AG-IC100 KN Electronic Universal Testing Machine (SHIMADZU, Japan).

UK using DMLS EOSINT-M270 machine supplied by EOS GmbH, Munich, Germany.Uniaxial compression tests were carried out using an AG-IC100 KN Electronic Universal Testing Machine (SHIMADZU, Japan). Experiments are performed at a constant loading rate of 0.75mm/min which equates to an axial strain rate of about 10- 3/s, according to ISO-13314-2009. The compression responses of tested specimens were recorded and used to calculate the stress train curves of each TMPS cellular structure.Fig 1 Experiment schematic of the uniaxial compression test2.2 Finite element analysis modelingThe finite element analysis was carried out with commercial software, DEFORM- 3D. A full size model of 25mm×25mm×12.5mm was used to investigate the mechanical properties. The TPMS cellular structure models were placed between two parallel plates, which are regarded as rigid materials.

The TPMS cellular structure models were placed between two parallel plates, which are regarded as rigid materials. The bottom plate remained stationary while the top plate moved downward at a constant speed (0.0125mm/s) to compact the specimen.Fig 2 FEA model of the uniaxial compression testIn this work, the Elastic-plastic model was used, and the flow stress at zero strain represents the yield stress for the material. The yield stress increases with the accumulated effective plastic strain. The strengthening mechanism was considered as Johnson-Cook (JC) model, which is commonly used in the dynamic response analysis of materials.Results and Discussion3.1 Finite element analysis results2321Failure under compression testing is highly related with the stress distribution on the surfaces of the specimens. In order to illustrate failure mechanisms of porous structures, it is essential to analyze stress distribution under compression.

In order to illustrate failure mechanisms of porous structures, it is essential to analyze stress distribution under compression. As mentioned above, the straight edges and sharp turns in lattice cellular structure will result in highly stress concentration[8]. Compared with the existing lattice cellular structures (BCC, BCCZ, FCC, diamond et al.), triply periodic minimal surface structures have smooth infinite surfaces and uniform curvature radius.In this work, Fig.3 (A) shows the Von Mises stress distribution of the Gyriod structure at 5% overall deformation. The plot shows that the stress distribution is similar in each unit cell and the stress on the surface of each strut is uniform, while the 45°struts concentrate more stress than the horizontal struts. As many researchers reported [7, 11], for both bulk and porous structure materials, failures occur along 45° direction, which is the biggest shear stress direction.

As many researchers reported [7, 11], for both bulk and porous structure materials, failures occur along 45° direction, which is the biggest shear stress direction. Figure 3 (B) shows the shear stress distribution of Gyriod structure at 5% overall deformation. In this figure, it is clear that the shear stress concentrates on the 45° struts, and the level of shear stress is highest at the center of the struts, which is because the strut center has the minimum cross-sectional area.ABCDFig.3 Stress distribution of the Gyriod cellular structure at 5% overall deformation: (A) Von Mises effective stress distribution,(B) 45°shear stress distribution,(C) Von Mises effective stress distribution on the vertical plane, and(D) 45°shear stress distribution on the vertical plane.Clearly, in the unaxial compressive test process, failures are easy to first occur at the 45° struts of the Gyriod structure, especially at the center of these struts.

Hence, the direction of fracture zones is approximately 45 degrees to the horizontal plane. Furthermore, to investigate the failure mechanism, the Von Mises stress and shear stress distributions of a vertical cross-section of the whole model are showed in Fig.3(C) and (D), respectively. Then, the stresses at different locations in the struts are observed and it is found that the stress level on the strut surface is higher than that inside of the strut.2322Therefore, plastic strain firstly occurs on the surface of the struts. Since the actual surfaces of the SLM-manufactured objects exist a large number of defects (such as adherent unmelted powder particles, micro-cracks and micro-pores)[10, 12, 13], macro cracks are easy to be initiated and propagate along these defects induced by high level stress.

So, it is more likely to produce cleavage fracture.3.2 Experimental resultsFigure 4 shows the actual failure position and fractured zone directions at the compression tests after 30% deformation. In this work, failure position mostly occurred at the center of the struts, while there were still some failures at the positions near the joint regions mainly due to the surface defects of these regions. Nearly all the failures happened on the inclined struts and few failures on the horizontal struts. The fractured zone directions in Fig. 4(B) and Fig. 4(C) are all about 45 degrees to the horizontal plane.The fracture morphologies of the compression sample are shown in Fig. 5. The smooth shell and river patterns appear in Fig. 5(A) being the characteristics of cleavage fracture. However the small and irregular dimples in Fig. 5(B) are also distributed around the cleavage fracture shells.

5(A) being the characteristics of cleavage fracture. However the small and irregular dimples in Fig. 5(B) are also distributed around the cleavage fracture shells. Therefore, the fracture mode of the Ti-6Al-4V Gyriod structure is a mixed style based on brittle fracture. Similar conclusion for the fracture mode of the Ti-6Al-4V BCC porous structure fabricated by SLM was made by[14].ABCFig.4 Observations from different perspectives after 30% deformation: (A) Fracture positions observed from the top view, (B) Fracture zone direction observed from the front view, and(C) Fracture zone direction observed from the back view.ABCFig.5 Fracture and surface morphologies of the test sample after 30% deformation: (A) Smooth river and shell patterns, (B) Small and irregular dimples, and(C) Surface2323crack initiation and propagationThe surface morphology of the SLM-made lattice structure in Fig. 6(C) shows the crack initiation and propagation.

6(C) shows the crack initiation and propagation. The fractured surface of the SLM-manufactured specimen is rough with many pores and unmelted powder particles. The cracks initiate at the defect positions and propagate perpendicularly to the surface, which is in agreement with the FEA analysis results mentioned above.ConclusionFailure of porous structures occurring in the uniaxial compression process is highly related with the stress distribution. In this paper, both finite element analysis and experimental methods are employed to investigate the failure mechanism of the SLM- manufactured Gyriod cellular structure under uniaxial compression, and stress distributions are also plotted in color contour.

Both the finite element analysis and experimental results show that the 45°struts concentrate more stress than the horizontal struts, and the level of shear stress is highest at the center of the struts, which results in the fact that the fracture zones all own 45° angle with horizon plane. Failures occur mostly at the center of the struts, which are opposite to the fracture mode of other porous structures having many straight edges and sharp turns.AcknowledgementsThe study was supported by the National Natural Science Foundation of China (Grant No. 51671091), the independent R&D subjects of Huazhong University of Science and Technology (Grant No. 2017JYCXJJ005).References[1] Thomsen, P., et al., Electron beam-melted, free-form-fabricated titanium alloy implants: Material surface characterization and early bone response in rabbits. J Biomed Mater Res B Appl Biomater, 2009. 90(1): p. 35-44.[2] Tancogne-Dejean, T., A.B. Spierings, and D.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFramework for CAD to Part of Large Scale Additive Manufacturing of Metal (LSAMM) in Arbitrary DirectionsJ.(cid:3)Logan McNeil*, William R. Hamel*, Joshua Penney*, Andrzej Nycz†, and Mark Noakes†Department of Mechanical, Aerospace, and Biomedical Engineering, University of Tennessee, Knoxville, TN 37916 † Manufacturing Demonstration Facility, Oak Ridge National Lab, 2370 Cherahala Blvd NTRC-2, Knoxville, 37932, TN, USAAbstractThe purpose of this research is to provide a framework for Large Scale Additive Metals Manufacturing (LSAMM) in arbitrary directions. Traditionally, slicing and path planning is done along the gravity aligned direction of a part, causing more complex geometrical shapes to have unsupported overhangs. The overhangs can be managed using a part positioner or a powder bed process.

The overhangs can be managed using a part positioner or a powder bed process. A different framework for slicing and building parts out of gravity alignment could improve current capabilities of LSAMM processes. The presented research focuses on segmenting more complex geometrical parts into gravity aligned (GA), non-gravity aligned (NGA), and transition segments to help generate toolpaths. Initial research of segment planning for complex geometrical shapes will be presented, as well as current results from builds completed at the University of Tennessee- Knoxville. The completed builds show that more consistent thermal evolution of a part based on the path sequence and torch angle results in successful builds.IntroductionMetal additive manufacturing technologies have had significant research advancements in the past 10 years. Additive Manufacturing (AM) is desirable for companies due to the ability to achieve complex geometry with the reduction in waste material.

Additive Manufacturing (AM) is desirable for companies due to the ability to achieve complex geometry with the reduction in waste material. However, issues have persisted with the strength and the production of near net shape parts [1–3]. AM techniques have become specialized to improve upon these issues by changing their material and power sources. If traditional manufacturing techniques can be leveraged for AM, the cost of building parts can be reduced. A summary of additive versus conventional manufacturing can be seen in Figure 1.Metal additive manufacturing describes a variety of different processes used to build metal parts. In general, two types of AM techniques persist in industry: Directed Energy Deposition (DED) and Powder Bed Fusion (PBF). PBF technologies were developed to give metal parts high resolution, but generally have a lower deposition rate [4]. Highly specialized cabinets are needed to control the layer height, air flow, and heat.

Highly specialized cabinets are needed to control the layer height, air flow, and heat. DED with powder feed technologies tried to integrate powder with a laser heat source, which allowed for faster deposition. However, these processes have proved more complex to implement the production of powder expensive. be can Electron Beam FreeForm Fabrication (EBF3) and Wire and Additive Arc Manufacturing (WAAM) are both DED processes with wire feed. The EBF3 process, while having a large deposition rate, requires a specialized vacuum cabinet.

The EBF3 process, while having a large deposition rate, requires a specialized vacuum cabinet. The WAAMFavor Conventional Manufacturing I Large Production Volumes Low Material Costs Easily Machined Materials Centralized ManufacturingI Favor Additive Manufacturing Lower Production Volumes High Material Cost High Machining Costs Capital Investment Logistics Cost Transportation Cost PrototypingandFigure 1 - Table from [4] that shows pros and cons of additive manufacturing.1126technique describes any arc process similar to welding that is used to deposit material. Current LSAAM focuses on using laser or arc sources to have higher deposition rates [5].

Current LSAAM focuses on using laser or arc sources to have higher deposition rates [5]. The WAAM process has been demonstrated for large scale structures like excavator arms [5], [6], but the full capabilities of WAAM have yet to be explored.Wire and Arc Additive Manufacturing (WAAM) WAAM is a desirable process since traditional manufacturing has relied on robotic WAAM systems to complete simple welding tasks for many years. Most manufacturing centers already have the basic manipulator and welding technology available and the cost for metal wire is relatively inexpensive compared to the powders for laser processes [5]. A setup for a typical WAAM cell can be seen in Figure 2.(a) Picture of the welding workstation in (b) CAD model of Kuka KR6-2 with sensor Dougherty 110.bracket attached.Figure 2 - Illustrates the welding cell and setup at the University of Tennessee.WAAM research has focused on building parts to near net shape [1], [7-10].

As a standard manufacturing technique, there has been significant research on weld pool shape [11], [12], thermal behavior [13], [14], and the microstructure of the resulting welds [6], [15]. Most of the current research on large scale WAAM has been shown in the GA orientation [ Typically, this is accomplished by using a welding system and part positioner to move the part to allow the torch to stay in the GA orientation [5], [10]. While demonstrated for large scale parts with large features, more complex geometry will require for the torch to complete deposition in NGA orientations.Traditional CAD to Part for AM One of the most attractive parts of AM is the CAD to Part methodology. Part designers are able to design a part in their Computer Aided Design (CAD) software of choice, upload the file to a Computer Aided Manufacturing (CAM) software for their machine and finally build the part.

A traditional design and build approach for AM can be seen in Figure 3.The traditional additive manufacturing CAD to Part process can be simplified into: Design, Slicing, Machine Code Generation, and Simulation [16]. The phases are interrelated, but have their own unique effects on the part plan. Some additive techniques use real time control to ensure build success, but they still rely heavily on the pre- part planning done in software programs like Insight or Cura. The goal of the CAD to Part process is to ensure that the part plan for the build maintains geometrical accuracy and strength of the final part. Design focuses on the part shape and the resulting build based on the additive technique. If a certain technique cannot complete the feature type, it must be redesigned keeping the processing parameters in mind [17]. Slicing typically focuses on two types of part representation: boundary or volumetric representation [18].

Slicing typically focuses on two types of part representation: boundary or volumetric representation [18]. Slicing the correctly represented 1127model allows the part planning software to generate the correct robotic moves to improve thermal evolution and shape filling for AM processes. The software then takes the slicing plan and generates the proper machine code to build the part using the specified AM process. The machine code is represented generally by a G-code, but interpretation into machine specific code can be done in-line or by the system controller. Finally, simulation can be done on slicing, machine code, or of other properties to ensure that the part plan can be achieved.

Finally, simulation can be done on slicing, machine code, or of other properties to ensure that the part plan can be achieved. Iterating through all of the steps should provide a plan with ideal slicing and machine code to achieve near net shape of the part.-------( I I TESSELLATION 1 I I\, ______ .,,,PART ORIENTATION~ - , ______ .,,S~P;O:T- - STRUCTURES\ GLICINGMACHINE TOOLPATHr-:-7 L:::_JCAD CAMExternal Hollowing lnfills Microstructures Frame structuresUniform Adaptive DirectVector RasterMat. Deposition Mat. SolidificationFigure 3 - Traditional CAD to Part Method from [16].Current MethodsThe first goal of this research is to extend the ability of LSAMM in arbitrary directions to make up for the lack of support material and to increase the amount of achievable geometrical complexity. Typically, LSAMM has been accomplished with the deposition being done aligned with gravity.

Typically, LSAMM has been accomplished with the deposition being done aligned with gravity. Researchers at the University of Tennessee designed an example complex part that contained the three different types of sections: GA, NGA, and transition segment. The CAD representation can be seen in Figure 4. The designed part will be used as an example to show how different pre planning techniques can be used to help integrate complex geometry.The current framework for building parts for complex geometry is outlined in Figure 5. Initially the CAD of the part is built in Solidworks. The CAD model will then be sliced and the sequencing of the part plan specified in MasterCam. MasterCam gives the ability within each layer to set travel speed, directions, and start/stop points. After the initial plan is made, the part plan is imported into Octopuz.

MasterCam gives the ability within each layer to set travel speed, directions, and start/stop points. After the initial plan is made, the part plan is imported into Octopuz. Octopuz is a robotic planning software where a model of the work-cell is used to assign welding settings and check the path for errors in the robot orientation. Octopuz then generates the robot code for the machine and is executed from the pendant. Following all of these steps allows the researchers to build up a part plan that has slicing, sequencing,Figure 4 - Example Part CAD and representation. Blue is GA, yellow is Transition, and green is NGA.1128Current LSAMM Preplanning Processand settings all specified by researchers.

Blue is GA, yellow is Transition, and green is NGA.1128Current LSAMM Preplanning Processand settings all specified by researchers. The current LSAMM preplanning process was verified by previous experimentation on simple geometry GA builds developed at the University of Tennessee-Knoxville.Slice Segments base~ on height and segmen typeIm lement Directions and Tool Parameter SettingsentSliceFigure 5 - The preplanning process that includes design, segmentation, slicing, settings, and simulation.Initial Complex Box Build The initial complex build was attempted from a part similar to the one seen in Figure 4. The purpose was to test the capability of the planning framework as well as to see the capabilities of the system to complete NGA segments with open loop control. The segments were planned according to the flow diagram seen in Figure 5.The initial box build part plan will contain three sections of planning: GA, NGA, and transition.

The segments were planned according to the flow diagram seen in Figure 5.The initial box build part plan will contain three sections of planning: GA, NGA, and transition. These representative sections make up the different types of segments for LSAMM. The blue section in Figure 6 represents the GA section where the deposition occurs with the torch in gravity alignment. The start/stop points are the most concerning sections and these can be mitigated by moving the torch to different parts along the path for the start/stop points. Alternating directions on each layer also allows the part to develop a more consistent thermal evolution between each layer. The sequencing of the GA previous was experimentation GA structures on this system.takenfromSegment Planning for GA, Transition, and NGA Sections - Initial PlanningLayer 1Layer 2Layer 1Layer 3Layer 1Layer 2onLayer 4The transition segment planning is also handled in a similar fashion.

As a simple plan of the transition zone, the researchers decided to keep the transition segment gravity aligned to ensure build success. The build results in a stair-step effect along the interface between theFigure 6 - Path Plan Outlining the GA, Transition, and NGA section. The green arrows indicate starts of layer while red indicates the end of a layer.1129NGA and transition segment. Since the segmentation is stair stepped, it allows the torch to remain in the GA position for the entire GA and transition segment. The simplified planning of this transition zone was used for initial feasibility tests, but future research will focus on more complex transition planning techniques.The final section seen on the far right in Figure 6 is the NGA section sequence. The sequencing follows the same order as the GA section, with the only difference being the torch angle. The torch angle for both the horizontal and vertical done perpendicular to the 35-degree NGA orientation.

The torch angle for both the horizontal and vertical done perpendicular to the 35-degree NGA orientation. The torch angle was done at this angle for the ease of the path plan program as well as a simple first proof of concept for the NGA sections. A diagram of the torch angles for the NGA can be seen in Figure 7. A layer height of 2mm was chosen for this build based from previous GA testing on the system. The settings for each of the following builds were not changed only the slicing and sequencing of the part were varied to help reach near net shape. Future research will focus on the effect that welding settings can play in the NGA sections.sections wasFigure 7 - Torch Angle for Horizontal and Vertical sections for initial complex part build NGA section.The resulting part can be seen in Figure 8. The GA and transition zone sections of this build worked well due to previous work with building other GA parts. The transition zone had a minor issue due to needing to skip a layer.

The GA and transition zone sections of this build worked well due to previous work with building other GA parts. The transition zone had a minor issue due to needing to skip a layer. This occurred because the part was growing more rapidly than the part plan had predicted. This occurred around layer 12 of the transition zone. Only three layers of the NGA section were able to be completed due to excessive drooping and issues with layer height. The omission of a layer in the transition zone also altered the total height of the part, thus causing it to deviate from the desired part plan. The decreased height at the top of the transition zone caused long arcing issues with the welding process that also contributed to the ending the build before completion.

The decreased height at the top of the transition zone caused long arcing issues with the welding process that also contributed to the ending the build before completion. In general, only the NGA section was unsuccessful, so future builds were designed to focus on improving the NGA sections using variable start/stop points, variable sequencing, and variable torch angles.Edge not completedFigure 8 - Results of Initial Complex Part Build. The build was successful up until the NGA sections.NGA ExperimentationAfter the initial complex part was unsuccessful, researchers decided to try and isolate the NGA section for further experimentation. The goal was to determine what best techniques could be used to help counteract the effect that gravity plays on the weld pool in NGA orientations. Initial work from [2] helped guide initial ideas on varying torch angle to help counteract gravity effects.

Initial work from [2] helped guide initial ideas on varying torch angle to help counteract gravity effects. Researchers also focused on defining a sequencing1130technique for the NGA section to help improve the ability of welding in NGA sections. Researchers decided to weld only the NGA sections at 45-degrees to test different ideas on improving NGA sections.Initial 45-Degree Test The first test was to see if rotating the torch would help counteract the gravity effects seen in the NGA sections. Initially, researchers modified the torch angle just slightly from parallel and attempted to build with a similar sequence to the NGA section seen in Figure 6. The torch angle plan can be seen in Figure 9. The torch angles were just slightly modified and the layers alternated direction and started and ended on the midpoint of the four sections.The results of the build can be seen from Figure 10.

The torch angles were just slightly modified and the layers alternated direction and started and ended on the midpoint of the four sections.The results of the build can be seen from Figure 10. The results were surprisingly successful, as varying the torch angle helped hold the weld pool into place. The programmed path was expected to build up to 30 layers, but the end result did not reach this due to the variability of welding uphill and downhill. The uphill sections were much harder to weld, and resulted in the bumpiness seen on the vertical sections in Figure 10. The horizontal sections had a much cleaner top geometry since they were held up by the torch angle during deposition. The bumpiness and variability of layer height contributed greatly to the part plan not predicting the exact height during the build.

The bumpiness and variability of layer height contributed greatly to the part plan not predicting the exact height during the build. However, in general the initial NGA section welded at 45 degrees was much more successful than the initial complex part NGAFigure 9 - Initial 45 Degree Torch Angles with similar layer sequence. The horizontal(left) and vertical(right) sections have the same torch angle variability.Therefore, it was decided from this result that a different layer sequence must be attempted that could eliminate uphill welding. Since the sequence was changed, there would have to be more stop/start points on the top layer to only weld downhill.

Since the sequence was changed, there would have to be more stop/start points on the top layer to only weld downhill. Varying the torch angle also helped the vertical and horizontal sections hold the weld pool against gravity and provided more stability for the arc in the NGA orientations.Downhill Welding Sequences After determining that varying only torch angle was not sufficient for producing consistent layers for NGA sections, researchers decided to modify the sequence to only have downhill welding sections in the vertical sections. The horizontal sections would have multiple start/stop points, and the effects of welding downhill to produce consistent layers could be identified. By combining a variable torch angle and a unique sequencing for NGA sections a consistent layer growth could be achieved for both the vertical and horizontal sections.Figure 10 - Initial 45-degree part results.

Note the bumpiness in the vertical sections while the horizontal sections had a much smoother surface.1131Layer 2Layer 2bFigure 11 - New two-start sequence. Green indicates the starting layer with red indicating the end of that layer. The layers are alternated for the entire build. The initial new sequencing is outlined within Figure 11. The sequencing required two start/stop points along the top layer so that both vertical sections could be welded downhill. Initially, only minimal layer heights were achieved due to the effects of the welding angle along with the new sequencing. Researchers decided to isolate the effects of the sequencing from the angle to help ensure that the new sequencing was effective in producing a closer near net shape than the old sequence. If more layers could be built than the previous NGA segment using a different sequence, it can be considered an improvement on the stability of the process to grow layer to layer.

If more layers could be built than the previous NGA segment using a different sequence, it can be considered an improvement on the stability of the process to grow layer to layer. The result of the new sequence perpendicular to the 45-degree plate in both the horizontal and vertical section can be seen in Figure 12.While more layers were achieved than the initial 45-degree part in Figure 10, the results of the layer growth are vastly different. The upper surface is much smoother to the eye in the two-start sequencing, but the layer growth was more consistent when the welding sequence included uphill and downhill welding in the vertical sections. The layers also seemed to build up on the bottom of the part, since the weld pool was always traveling downhill.

The layers also seemed to build up on the bottom of the part, since the weld pool was always traveling downhill. On the horizontal sections, the start/stop points could easily be identified due to a variation of heat and mass input from the arc ramp up and down for beginning and ending layers.Figure 12 - Side and Top View for new sequence 45-degree build.While the results were a mixed success, researchers considered that a combination of proper torch angles in the vertical and horizontal sections alongside a new sequencing should improve the ability to build new NGA sections. Note for all of these builds the process parameters were not changed; only sequencing and torch angle were improved. All of these builds were completed open loop, without any real time control to keep layer height consistent.

All of these builds were completed open loop, without any real time control to keep layer height consistent. Ideally a robust LSAMM system would have a combination of pre planning and real time control strategies to help maintain a consistent layer height and resultant layer geometry.New Complex Part Build Researchers were encouraged from the results of the 45-degree builds to attempt the complex part again with a new torch angle and sequencing. The same process was followed from the first complex build by building the path plan through MasterCam and Octopuz to allow modification and simulation of the machine code for the 1132robot. The design of the new sequencing can be seen in Figure 13. The sequencing of the NGA section was modified to have 5 start/stop points along the horizontal sections to help mitigate the effects of the humps seen in Figure 12.

The sequencing of the NGA section was modified to have 5 start/stop points along the horizontal sections to help mitigate the effects of the humps seen in Figure 12. All vertical sections were welded downhill to help improve top surface quality and to have a more consistent layer height for the build. The sequencing of the GA and Transition segments were not changed from the initial complex part build.Layer laLayer lbSegment Plann ing for GA, Transition, and NGA SectionsLayer 1ayer 2aLaye r 2Layer 3aFigure 13 - New Sequencing of the Complex Part. Note the new 5-layer sequence of the NGA Section.Layer 3bThe torch angles for the new complex part can be seen in Figure 14. The angles were torch slightly different in both the and vertical horizontal sections. the horizontal Since sections worked relatively well at anyLayer 3angle, they were only put 5 degrees off the 45-degree plate.

the horizontal Since sections worked relatively well at anyLayer 3angle, they were only put 5 degrees off the 45-degree plate. The vertical sections had a steeper torch angle to help mitigate the downhill welding pooling effect at the bottom of the section, in an effort to get similar growth in the vertical and horizontal sections. The welding settings were the same for this part as were on the initial complex part build, as to keep consistent results.The results of the build can be seen in Figure 15. The build was successful and completed all 30 programmed layers of the NGA section. The new sequencing and torch angles proved successful to build the NGA sections. It is determined now that it is more than feasible that through a combination of new sequencing, angles, and power settings that NGA segments can be implemented into LSAMM for more complex geometries.

The success ofFigure 14 - Torch angles for the horizontal and vertical sections of the new complex geometry sequence.Figure 15 - Results of the new complex geometry with a new sequencing for the NGA sections.1133these builds and experimentations can be used to help modify the existing framework for LSAMM planning which should help part planners reach a better near net shape.Conclusions and Future WorkFrom the current research results, researchers have proved that by modifying just the sequencing and torch angle in the path planning section will help improve near net shape of complex geometries with NGA sections. The effects of variable bead height, more robust real time weld pool control, and new types of sequencing and slicing strategies need to be considered for future work. Future research will focus on extending the ability of LSAMM systems to build complex parts with integrated NGA sections.

Future research will focus on extending the ability of LSAMM systems to build complex parts with integrated NGA sections. Commercially available slicers do not typically have this type of capability so tools to build integrated complex geometries need to be developed to improve research in this area. If these tools can be developed and integrated with experimental results from NGA welding segments, novel approaches to planning and programming LSAMM parts using WAAM can be considered. A deeper dive into horizontal versus vertical sections during LSAMM using WAAM has begun, with initial results that can be seen in Figure 16. The walls in Figure 16 were produced with a simple sequence and perpendicular torch angle; but the real effect of colder welding settings can be seen to produce overhangs up to 105 degrees – if not further.

Framework for Large Scale Fused Pellet Modeling (FPM) by An Industry RobotZhiyuan Wang, Renwei Liu, Todd Sparks, and Frank LiouDepartment of Mechanical & Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO, 65409AbstractFused pellet modeling (FPM) is an important method in additive manufacturing technology, where granular material is used instead of filaments. In FPM, prototypes are constructed by the sequential deposition of material layers. As the size of the part increases, the problem of long build times and part deformation becomes critical. In this paper, methods for eliminating the void density during deposition and accuracy control principles for large scale FPM processes are studied. By analyzing the ab initio principles of this process, a mini extruder with variable pitch and progressive diameter screw for the large scale fused deposition is proposed.

By analyzing the ab initio principles of this process, a mini extruder with variable pitch and progressive diameter screw for the large scale fused deposition is proposed. Based on polymer extrusion theory and non-Newtonian fluid properties, each of the design parameters are analyzed, such as the length of different function sections of screw, die shape of extruder nozzle, and the material properties. According to these analysis results, an extrusion process simulation for controlling the filament shape is carried out with multi-physics modeling software and proved the FPM could increase the building efficiency and deposition quality for large size parts.IntroductionLayered manufacturing with extruded material is one of the most promising rapid prototyping techniques to demonstrated novel design ideas and reduce the product development cycle.

This process fabricates prototypes by extruding the material in semi-fluid status through a heated nozzle in a prescribed pattern onto a platform. Various types of material could be applied in this process including polymer, cement, plaster, and wax.The deposition material should be extruded continuously, stably and under constant temperature during the layered manufacturing process. “Continuously” means there should be no interruption of extrusion when the nozzle is scanning the deposition path; “Stably” involves the stable extrusion amount and accurate geometry of the semi-molten material; “Thermostatic” is to ensure the bonding quality between deposition tracks.The fused deposition modeling (FDM) developed by Stratasys Inc., has been a leading rapid prototyping technology, which involves layer by layer deposition of extruded material through a nozzle using feedstock filaments from a spool [1].

The material feeding process of FDM realized by two friction wheels rotate reversely to push the filament into a heated nozzle (Figure 1). Because of the advantages of simple structure and easy control, it has been widely used in most of the fused deposition systems. But the weak points of this method are also very obvious, the extrusion force is limited by the surface compressive strength of filament and the contact area between friction wheels and filament. Insufficient friction will cause slip feeding, and too much compressive force applied on the filament might be break it off, both of these will affect the extrusion quality. To shift from rapid prototype to agile fabrication by broadening thematerial selection and improve the properties of fused deposition modeling, Anna Bellini et al [2] presented a novel extrusion system which mount on a high precision positioning system and fed with granulated material.

Although this screw extrusion system could perform better than other filament based extrusion systems, but the normal single screw is still not the best suitable structure for the fused material extrusion process.Figure 1 Schematic of filament based extrusion systemsAs the size of the part increases, the problem of long build times and part deformation becomes critical. Currently, most FDM machine’s extrusion amount is between 12-17 kg/hour, it will take several days to build a big size part. Inner stresses resulting from the contraction of deposition fibers within one layer can affect the prototype size precision, bring about prototype deformation, including warp and inner-layer delaminating or cracking. Similar quality issues widely exist in other rapid prototyping process. Many researchers have investigated the inner stresses. Tian-Ming Wang et al. [3] constructed a mathematical model of the prototype warp deformation to analyze each of the influencing factors. Jayanthi et al.

Tian-Ming Wang et al. [3] constructed a mathematical model of the prototype warp deformation to analyze each of the influencing factors. Jayanthi et al. [4] discussed how the scanning pattern of the laser in stereolithography (SLA) influences the resulting deflection of the part. Céline Bellehumeur [5] study the dynamics of bond formation between polymer filaments with thermal analysis and sintering experiments under different conditions and the same group performed a more in-depth study on the mechanisms that control the bond formation under different process conditions [6].To avoid most defects of material extrusion steps in a filament-based system, a mini screw extrusion system has been proposed and developed in this paper, and especially designed for building large size object with layered manufacturing technique.

The new set-up, called fused pellet modeling (FPM) system and consisting of a mini extruder mounted on an industry robot arm, operates using bulk material in granulated form. This configuration opens up opportunities to use wider range of materials, meanwhile the robot arm making the fused pellet modeling process more flexible and become a viable alternative rapid prototyping for large size part.System specificationScrew extrusion has been widely adopted in the polymer manufacturing industry because of its extremely excellent attribute for processing. As the plastic moves along the screw, it melts by the following mechanism. Initially a thin film of molten material is formed at the barrel walls. As the screw rotates it scrapes this film off and molten plastic moves down the front face of the screw flight. When it reaches the core of the screw it sweeps up again, setting up a rotarymovement in front of the leading edge of the screw flight.

When it reaches the core of the screw it sweeps up again, setting up a rotarymovement in front of the leading edge of the screw flight. Initially the screw flight contains solid granules but these tend to be swept into the molten pool by the rotary movement. As the screw rotates, the material passes further along the barrel and more and more solid material is swept into the molten pool until eventually only melted material exists between the screw flights.Applying the screw extrusion method into fused deposition process is not simply copy, since the nozzle needs to keep certain scanning speed during the deposition process, it requires that the weight and volume of extrusion unit should be within a certain range. The key component in the system is extrusion screw.

The key component in the system is extrusion screw. Some researchers have tried to use common screw for extrusion, it could work for several kinds of material, like ceramic and plaster, but this structure cannot provide uniform mixing and enough back pressure for polymer material in fused deposition process. To solve these problems, a mini screw with variable pitch and progressive diameter is designed and manufactured for fused deposition purpose.In combination with the barrel, the purpose of the screw is to convert solid material to the melted status, and finally pump material to the die in an efficient manner. To obtain a good extrusion performance, the screw for FPM system should be designed as three different function sections: feed section, transition section and metering section.

To obtain a good extrusion performance, the screw for FPM system should be designed as three different function sections: feed section, transition section and metering section. The feed section is the first element of the screw where the polymer pallets is introduced to the screw, the wide pitch P1 and small diameter D1 the could provide the most feeding volume and frication force for pushing material to move forward. The transition section (or compression section) is where most of the melting of polymer takes place. This is portion of the screw that “transition” from the feed depth to the metering depth and where work is done on the resin causing melting to occur. In this section of screw, the root of the screw gradually becomes shallower forcing the material towards the wall of heated barrel where the melting takes place.

In this section of screw, the root of the screw gradually becomes shallower forcing the material towards the wall of heated barrel where the melting takes place. The last but not least is the metering section or pumping section of the screw is where the melting of the polymer is completed and pumping to overcome the head pressure takes place.This mini pallets extruder is mounted on an industry robot as shown in figure 2(a), it has a 4.1 m2 (cross-section area) operating area and 300⁰ rotation range for the base motor (figure 2(b)), which could provide much bigger deposition working envelop than other current fused deposition system.

The 6-axis movement mechanism makes the deposition process more flexible to build the model with complex features.(a) Assembly model of FPM system (b) Robot working envelopFigure 2 Assembly model of FPM system on the industry robot and working envelopAccuracy control of extrusion processThe void density caused by gap and overlap between deposition tracks is inevitable defects of filament based FDM system, it also will result the part deformation and failure deposition. To analyze the effect of void density in deposition process, L. Li [7] proposed a method to calculate it theoretically.

To analyze the effect of void density in deposition process, L. Li [7] proposed a method to calculate it theoretically. In an orthotropic laminate, in the plane normal to filaments or axis d, define1r as the area void density, g as the gap size.When g =0, there is no gaps among filaments, the ideal cross section is shown in Figure 3, the cross section shape of filament is an ellipse with a and b as the idealized lengths of semi- major axis.Figure 3 Ideal cross section of deposited filamentIn the actual process, the top and bottom surface of filament would flatten when deposited onto previous layers, this is resulted from the vertical force of extrusion and it is in semi-molten state. Therefore the modified calculation model is shown in Figure 4(a) when considering the flattening effect with d that can be measured experimentally.

Assuming that the total cross sectional area stays unchanged, the relation between area Q and P should be:Q  2P (1) Assuming the d is a small amount, the flatten area has the same length as the long semi-axis of filaments ellipse, thus the void density can be calculated using the following equation:pab 2a(2b d) The real cross section for this situation also indicated in the microscopic photo of Figurer1  1(2)4(b).(a) Schematic diagram of deposited filament with flattening effect(b) Microscopic cross section of deposited filamentFigure 4 Deposited filament with flattening effect and photo of microscopic cross sectionWhen g  ±D , which means there is a gap or overlap between deposition tracks, shownin figure 5 and figure 6, and cross section photo taken under microscopic, respectively.r1  1pab (2a ± D)(2b d)(3)(a) Schematic diagram of deposited filament with positive gap(b) Microscopic cross section of deposited filamentFigure 5 Deposited filament with positive gap and photo of microscopic cross section(a) Schematic diagram of deposited filament with negative gap(b) Microscopic cross section of deposited filamentFigure 6 Deposited filament with negative gap and photo of microscopic cross sectionThe smaller of area void density means better deposition quality, it could be found that increase d and use minus gap setting will decrease the 1r , but this requires more deposition layers and denser deposition tracks, it will result longer building time and d also has a limitation because the structure of nozzle.The polymer material changed three different physical states sequentially from it been fed into the barrel to be extruded out through the nozzle: glass state, high elastic state and viscous flow state.

Molten polymer has a high viscosity and significant elasticity, it should be considered as a kind of viscoelastic material and exhibit non-Newtonian fluid properties. One important feature of non-Newtonian fluid is Barus effect [8], which is an expansion phenomenon of a non- Newtonian fluid when it emerges from a nozzle to an open space so that the diameter of the emerging stream can be several times the nozzle diameter and the expansion ratio is varied from the edges to corners. When the molten polymer extruded out from the nozzle, it will expend along the edge, the longer of the edge the more obvious of expansion. This will result the filament shape of extruded material is different from the shape of nozzle (figure 7).

This will result the filament shape of extruded material is different from the shape of nozzle (figure 7). Therefore, if the expected filament shape is a rectangle, the shape of nozzle should be similar with third one in figure 7.Figure 7 Barus effect of different shape nozzleBy controlling the parameters of extrusion process and the shape the nozzle, a rectangle shape of filament as shown in figure 8 could obtained for decreasing the void density and enhance the deposition quality.(a) (b)Figure 8 Model of nozzle for extruding rectangle filament and ideal deposition resultsExtrusion process modeling and discussion of simulation resultsFor non-Newtonian flow, denotes the viscosity (kg/(m·s)), u the velocity (m/s), r the density of the fluid (kg/m3) and p the pressure (Pa).

The equations to solve are the momentum and continuity equations.ru   t       (u(uT ) )r u    up0(4)0   (5)In the Carreau model, the viscosity depends on the shear rate, , which for anaxisymmetric model in cylindrical coordinates is defined according to Equation 2:1 2  (2 ) u r2u 2(rv r)2v (2 ) z24(u r)2  (6)The viscosity is given by     0( )[1 (2 ) ]( n 21)(7)where  is the infinite shear rate viscosity,0 is the zero shear rate viscosity,  is aparameter with units of time, and n is a dimensionless parameter.The section view of nozzle in this FPM system is shown as Figure 9(a).

To analysis the flow behavior of molten polymer material in COMSOL multi physics, an inside volume model of the nozzle has been built as shown in Figure 9(b).(a) (b)Figure 9 Cross section view of nozzle and the inside volume modelTo analysis the Barus effect of the extruded molten polymer, three different shapes nozzle are designed and the inside volume models of the nozzle are also modeled respectively (Figure 10).

Shape-1 represents the original rectangle shape die, the length of edge is 0.5 inch; Shape-2 represents the modified shape die, the edge contracted inward obviously, the radius of arc is equal to the length of the rectangle and the arcs are tangent to each other; Shape-3 represents the improved shape die, the edge contracted inward slightly, the radius of the arc is 0.7 inch, the radius of fillet at each corner is 0.01 inch.Figure 10 Cross section sketch and 3d model of three different shapes nozzleThe inside volume model represents the material which is extruding out from barrel through the die of nozzle. The flow behavior of melted material follows the above non- Newtonian flow equations. Because of the symmetry geometry feature and for the computation efficiency, the model could be simplified as single quadrant and imported into COMSOL multi physics.

Because of the symmetry geometry feature and for the computation efficiency, the model could be simplified as single quadrant and imported into COMSOL multi physics. The material subdomain settings for low density polyethylene (LDPE) at around 220 degree centigrade [9] are shown in table 1.Viscosity model typeCarreau modelDensity of melt LDPE (rho)743 kg/m3Zero shear rate viscosity1437.4Pa.sModel parameter (n)0.39Model parameter ()0.015sTable 1 Subdomain settings for LDPEWhen non-Newtonian flow comes out from the nozzle with certain speed and back pressure, the inner pressure of the flow will force the material expend in the open space, this the macro phenomenon of Barus effect, which will result the extruded material to swell and comes out with oval shape or other unexpected shape, this phenomenon could be reflected in the cross section pressure plot of simulation results.

Figure 11 shows the meshed inside single quadrant volume models of three different shape nozzle and the cross section pressure distribution plot at the outlet of nozzle, respectively.As shown in figure 11(a), the pressure along rectangle edge is higher than the pressure at the corner section, the filament surface will expend to a round similar shape, figure 11(b) shows the modified shape, which shrank too much for each edge and the pressure is concentrated at the corner because of the stress cusp effect. The results illustrate that simulation correspond the non- Newtonian flow’s characteristic and these two design cannot achieve the object of extruding rectangle shape filament, because the molten polymer cannot expend averagely.This problem can be solved by optimizing the die geometry and trying to even the stresses at the corners.

Figure 11(c) shows the pressure distribution plot of the improved shape die, the pressure at the corner areas is smaller than the edges area and the pressure distribution is nearly uniform and symmetrical, the stress cusp effect could also be eliminate by adding a fillet feature. The comparison between these three different nozzles shows that filament extrusion quality could be controlled with optimization of the die shape. Moreover, the rectangle shape filament will improve the deposition accuracy and efficiency for large scale objects.(a)(b)(c)Figure 11 Meshed single quadrant inside volume models and the cross section pressure distribution plot at the outlet of nozzleConclusions and Future WorkFilament based fused extrusion deposition has been innovated and developed for a long period.

It has the advantage of simple system structure and economical in cost, but the low extrusion speed and enclosed workspace limited applying this technique into building large size prototype. To solve these problem, a fused pellet modeling (FPM) method is proposed in this paper. A screw with variable pitch and progressive diameter is designed for providing sufficient extrusion material at high speed, also with a certain back pressure; By analyzing the cause of void density of fused deposition and based on Carreau model, the die shape was optimized for LDPE material in COMSOL.

For the next step, this project will investigate the coupling affection of multi-physic fields (including hydromechanics, thermodynamics and phase change) on the deposition process of fused pellet modeling (FPM), to achieve building high quality large size prototype and with efficiency.AcknowledgementThis research was supported by Laser Aided Manufacturing Processes (LAMP) Laboratory at Missouri University of Science and Technology. Their support is greatly appreciated. The author would like to thanks Maxwell Mulholland and Huixu Deng from Department of Mechanical & Aerospace Engineering at Missouri S&T for their valuable contribution to the project. Also all people attached directly or indirectly to the project.References[1] Masood S H. Intelligent rapid prototyping with fused deposition modelling[J]. Rapid Prototyping Journal, 1996, 2(1): 24-33. [2] Bellini A, Shor L, Guceri S I. New developments in fused deposition modeling of ceramics[J].

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA Framework for the Design of Biomimetic Cellular Materials for Additive Manufacturing T. McNulty*, D. Bhate†, A. Zhang†, M. A. Kiser*, L. Ferry**, A. Suder†, S. Bhattacharya#, P. Boradkar* *The Biomimicry Center, Arizona State University; †Phoenix Analysis & Design Technologies, Inc.; **School of Mathematical and Natural Sciences, Arizona State University; #Autodesk Inc.AbstractCellular materials such as honeycombs and lattices are an important area of research in Additive Manufacturing due to their ability to improve functionality and performance. While there are several design choices when selecting a unit cell, it is not always apparent what the optimum cellular design for a particular application is.

While there are several design choices when selecting a unit cell, it is not always apparent what the optimum cellular design for a particular application is. This becomes particularly challenging when seeking an optimal design for more than one function, or when the design needs to transition spatially between different functions. Nature abounds with examples of cellular materials that are able to achieve multifunctionality, but designers lack the ability to translate the underlying principles in these examples to their design tools. In this work, we propose a framework to bridge the gap between nature and designer.

In this work, we propose a framework to bridge the gap between nature and designer. We present a classification of natural cellular materials based on their structure and function, and relate them in a manner amenable for use in guiding design for Additive Manufacturing.IntroductionCellular materials are a particularly exciting research frontier in Additive Manufacturing (AM) due to their ability to improve performance in a wide range of applications, while minimizing material usage and structural mass. Some of these applications are classified in Figure 1. Since most cellular materials engineers work with tend to be composed of synthetic materials, it is easy to forget that humans have been using natural cellular materials like wood and cork for millennia. The first documented observations of cells in materials were conducted by Robert Hooke in 1665, who concluded that these cells were responsible for the curious behavior of the material (1).

The first documented observations of cells in materials were conducted by Robert Hooke in 1665, who concluded that these cells were responsible for the curious behavior of the material (1). Since then, many engineering structures such as honeycomb sandwich panels, metal foams and architectural domes, owe their designs, at least in part, to nature’s inspiration.Nature is full of examples of cellular materials. 3.8 billion years of evolution have resulted in cellular geometries that conserve material usage and enhance performance under the different conditions where life thrives (2). A wide range of natural cellular materials have been studied and their functions postulated, which are summarized in Figure 2.

A wide range of natural cellular materials have been studied and their functions postulated, which are summarized in Figure 2. The similarity between figures 1 and 2 is evidence of the strong correlation between functional benefit obtained through cellular materials in engineering applications and in nature.2188StructuralThermalFluidBiologicalWeight ReductionHeat ExchangersCatalyst CarriersBone IntegrationSandwich CoresFlame ArrestersPackagingCell growthStrain IsolationHeat ShieldsBuoyancyBiomimeticVibration ControlFigure 1. Application areas for cellular structures in engineering solutions that can leverage their special properties to enhance overall functional performance, adapted from (3)Figure 2.

Application areas for cellular structures in engineering solutions that can leverage their special properties to enhance overall functional performance, adapted from (3)Figure 2. Classification of functions of cellular structures in nature, adapted from (1)2189Recent advances in software have made it feasible to design structures that incorporate intricate cellular structures and AM technologies have enabled these designs to be realized. However, it is not always apparent to the designer which cellular geometry is ideal for a specific function. This becomes particularly challenging when designing for multiple functional objectives. It is this challenge that we seek to address by using a biomimetic approach to help a designer select an appropriate cellular structure for a specific function and integrate it into a larger structure.

We propose a four-level framework to integrate this biomimetic approach into an overall design and manufacturing strategy, as shown in Figure 3.Figure 3. Proposed framework for implementing a biomimetic approach in the design and manufacturing of parts with cellular materialsIn this paper, we seek to develop the top level of this framework. We first review the methodology we used to discover natural models (i.e. species in nature), and how this selection was narrowed in scope. Next, we present a set of candidate natural cellular design structures that we discovered, followed by a classification of these structures and the load-bearing functions that are of relevance to the design of engineering structures. Finally, we discuss limitations of this approach.MethodologyWhile the use of biomimicry to address problems and uncover opportunities is not new, it is only more recently that a methodology for applying the subject systematically has emerged (2).

The methodology can be used in two different ways: in the first approach, the solution to a specific challenge is sought in nature (“Challenge to Biology”). In the second approach, a study of the2190biological organism(s) leads the way to inspiring designs (“Biology to Design”). For this work, we used the first approach, but with the specific intent of finding examples of cellular materials in nature that manage structural forces effectively. With this approach, multiple design principles will emerge together to inform the design space. For our purposes, we define ‘cellular’ materials as those materials with a specific, repeating structure that forms the basis of the material itself. The specific steps we used in this work are as follows:1. The “Challenge to Biology” approach enables the discovery of natural models which are relevant to our functional needs within contextual considerations.

The “Challenge to Biology” approach enables the discovery of natural models which are relevant to our functional needs within contextual considerations. Through a wide survey of the available literature, we identified approximately 70 different natural models that were constructed at least partially of cellular materials, ranging from the well-known bee’s honeycomb to the Venus’ flower basket, a sea sponge that has a crack-arresting hierarchical lattice structure.2. Having identified these models, we studied and vetted our selection in the literature against the function and context of interest to establish the nature of the cellular material (design strategy), as well as evidence of a functional basis for the structure in question (i.e., the function of the material was for providing support to the structure).

In some cases this was a hypothesis bolstered with circumstantial observations, in others the structure’s functionality was validated through experimental or numerical techniques. Some natural models were discarded from consideration since the causality between structure and function could not be robustly identified.3. Finally, we narrowed the scope to examining how these specific cellular design strategies responded to, and managed, imposed loads.

Finally, we narrowed the scope to examining how these specific cellular design strategies responded to, and managed, imposed loads. This is a non-trivial aspect of the work since many natural cellular models have arisen (over evolutionary time) for more than one function, and a sub-optimal solution for managing forces may exist a trade-off for another function, such as buoyancy or thermal management.Natural Cellular Design StrategiesOur study of cellular materials in nature resulted in the identification of 11 models for which we identified adequate evidence of a causal link between observed structure and function, and further that this function was of a structural (load-bearing) nature. In our study, it emerged that cellular materials in nature could be broadly classified as being used to either fill space (3D), develop surfaces (2D) for protection, or use cellular design in cylindrical structures. These three categories align well with engineering structure design.

These three categories align well with engineering structure design. We discuss these natural models for each of these three categories in turn.1. 3D Space-Filling StructuresThe most significant driver for the use of cellular materials with AM is their potential to reduce material usage and mass – this is critical for aerospace applications but also relevant for the wider transportation sector as well as for the biomedical and construction fields. Mass reduction is most effective when it is applied to as much of the structure in question, and is thus relevant for designing 3D space-filling structures.2191Nature often seeks to minimize costly material usage, which has been described as one of “life’s principles” (2). The specific strategy for minimizing mass for a particular organism is a function of the specific loading conditions the structure is aiming to withstand, which in turn is influenced by the context the structure is experiencing.

ii. iii.Honeycomb: public domain, no attribution available or needed Venus’ Flower Basket: NOAA Okeanos Explorer Program, Gulf of Mexico 2012 Expedition Toucan: By Nicolas Billebault, via Wikimedia Commons2. Surface (2D) StructuresSeveral engineering structures involve enclosing an entity in an external surface. These span a range of scales from spacecraft and submersible hulls to aerodynamic skinsuits for cyclists. In keeping with the overall scope of this paper, we limit ourselves here to the discussion of surface structures that excel at managing mechanical loads and have identified a few organisms in Table 2 that achieve this through different strategies.2192Table 2.

As shown in Figure 4, these has led to a range of strategies for adding cellular designs within cylindrical structures to achieve these benefits.Figure 4. Design strategies for improving bending, ovalization and buckling resistance in cylindrical structures found in animal quills and plant stems, adapted from (20)2194Biomimetic Cellular Materials DesignThe main objective of this work is to propose a framework for the design of biomimetic cellular materials that can be integrated into the larger design and manufacturing system proposed in Figure 3. Towards this end, we need to identify the underlying patterns emerging from our biological models by using classifications that allow us to explore both the design options and the functional requirements. Finally, we need a framework that associates design and function to guide selection of a specific cellular design for a larger component. We deal with each of these three aspects in turn.1.

Finally, we need a framework that associates design and function to guide selection of a specific cellular design for a larger component. We deal with each of these three aspects in turn.1. Classification of Design OptionsThere are several ways to classify cellular materials (22; 23; 3). In the context of the present work, we take our cues from nature and propose a separation based primarily upon the use of the material for either 2D (surface) or 3D (space-filling) purposes, while also identifying strategies that cut across both approaches. This classification is compiled in Figure 5, with self-explanatory representations of the associated geometry.Figure 5. Proposed classification for cellular material designs, grouped in terms of 3D and 2D and cross- cutting concepts that apply to all cellular designs21952. Classification of Functions (Load-bearing)Natural and manmade structures both perform one, or more, functions in a larger system.

Classification of Functions (Load-bearing)Natural and manmade structures both perform one, or more, functions in a larger system. More fundamentally, however, a structure defines form and must have properties that enable it to retain this form while meeting functional requirements in its environment for the duration of its useful life. In engineering structures, the most important properties of interest are stiffness (or rigidity; the ability to resist permanent deformation), strength (the ability to resist buckling) and toughness (the ability to resist fracture). These properties are defined in context of the conditions experienced by the structure. With regard to the nature of loads, these are typically classified as compression, tension, torsion, bending and shear, or some combination of these (24).

With regard to the nature of loads, these are typically classified as compression, tension, torsion, bending and shear, or some combination of these (24). Further, these loads may be applied in one or many directions (such as uniaxial, biaxial and hydrostatic) and with different durations (such as varying strain rates, fatigue and vibration). Thus, the loading condition can be adequately described by specifying three pieces of information (see Figure 6):1. Loading type 2. Loading direction, and 3. Period of application(a)(b)(c) Figure 6. Proposed classification for loading conditions: (a) Loading type, (b) loading direction and (c) period of application21963. Structure-Function RelationshipsWith the classification established for both cellular material design and load-bearing functions, the final step in developing a usable framework is to relate the two.

In Table 4, we relate the function to the structure, which forms the basis of a design recommendation which we make limited to the organisms we identified in Tables 1-3.Table 4.

We have attempted to synthesize principles from several different natural models and while we have demonstrated that much is to be gained from such an approach, it is important to stress a few considerations that need addressing prior to implementation in design and manufacturing.1. Uniquely Biological ConsiderationsA key challenge faced by living organisms is the need to grow, and in several cases, the maintenance of function during growth (24). It is feasible that a specific structure is sub-optimal for a specific engineering application since it was constrained by this need. Another challenge with the biomimetic method is the de-confounding of the functional basis for structure since natural structures are often compromises for more than one (sometimes competing) function(s), and identification of the “weights” of optimization for these objectives is non-trivial.

In some cases, the effects of evolutionary baggage also result in sub-optimal design, where features are retained through evolution despite their inadequacies since they either have low inherent costs, or it was less expensive to make a particular compromise: the giraffe’s convoluted Recurrent Laryngeal Nerve (RLN) is one such example (26). This highlights the importance of extracting underlying principles, as we have sought to do here, rather than emphasizing one-off discoveries.2. Additive Manufacturing ConstraintsWhile AM processes greatly expand the possibilities for manufacturing parts with cellular materials, they do have additional constraints (i.e., ‘design rules’) that limit the selection of cellular material designs.

These constraints include:Smallest feature size attainable (typically in the hundreds of microns for most production- scale machines)Largest overhang in lattice-like cellular materials • Need for support in most processes, which constrains the cellular design to be self- supportingNeed for removing trapped powder, in powder bed processes, constraining the smallest size of voids possible in 3D structuresThese constraints need to be fully comprehended prior to the selection of a cellular material design.3. Hierarchy, Gradients and Multi-MaterialsCellular materials are just one of many design strategies found in nature (23). Cellular materials themselves consist of multiple materials such as the examples of the abalone shell and mantis shrimp club discussed previously, which currently still poses a challenge for functional part additive manufacturing.

Additionally, cellular materials often have functional gradients or a hierarchy of structure at varying scales, several of which may be below the ability of AM processes to resolve. Nonetheless, as AM technologies improve, it is likely that an increasing number of biomimetic designs are rendered feasible.2198ConclusionsIn this paper, we have demonstrated how a biomimetic approach can be used to enable the selection of cellular material designs for structural load-bearing applications.

Our work has led us to draw the following conclusions:Employing the biomimicry methodology of scoping and discovering complimentary models allows for a “Challenge to Biology” approach to identifying biological structures of interestNature tends to employ cellular materials in structural applications in three categories that align well with engineering design thinking: space-filling (3D), surface (2D) or cylindrical structuresNatural models tend to demonstrate hierarchy and the use of multiple materials, along with cellular material designs to achieve a specific performanceDefining a classification for cellular material structure and load-conditions, as we have done here, enables the formulation of a structure-function relationship that is the first step to developing an integrated framework for designing biomimetic cellular structuresIn ongoing and future work, we are seeking to use computational simulation to validate these biomimetic design concepts, which is a key step in developing a quantifiable correlation between structure and performance.

viewed PaperA FRAMEWORK WITH EXAMPLES FOR PRINTING THERMOSETTING POLYMERS USING LASER POWDER BED FUSION ADDITIVE MANUFACTURINGC. A. Chatham1 and A. L. Washington II21 Advanced Engineering Division, Savannah River National Laboratory, Aiken, SC 29808 2 Advanced and Energy Materials, Savannah River National Laboratory, Aiken, SC 29808AbstractThermoset polymers possess traits arising from their covalently crosslinked network structure that are distinct from thermoplastic polymers. These traits can manifest as relative increases in the performance properties of chemical resistance, stiffness, and strength particularly at elevated temperature. Although these performance properties are desired from additively manufactured parts, there are few engineering grade thermosetting polymers commercially available for any additive manufacturing (AM) fabrication method.

This work describes some of the challenges when processing and formulating thermosetting powder feedstocks for the laser powder bed fusion (L-PBF) mode of AM. The thermal curing properties of three different commercially available thermosetting polymers made via L-PBF are compared. Surface temperature profiles collected during printing are used to predict the extent of crosslinking through a rudimentary isoconversional model and are compared against post-print measurements of residual cure.IntroductionMany industries continue to increase their implementation of additive manufacturing (AM) technologies as layerwise fabrication transitions from primarily prototyping applications towards manufacturing of end-use parts. Of the seven technology categories identified by ISO/ASTM 52900 [1], powder bed fusion (PBF) is often championed for end-use production due to its batch- style fabrication, large build size, and low anisotropy compared to other AM processes [2, 3].

Additionally, the materials commercially available for PBF currently include relevant engineering grade polymers, like nylon and PEKK [4]. However, no one family of materials is adequate for every application. Certain applications demand a stiff mechanical performance even at elevated temperatures and in the presence of harsh solvents. Traditionally, these performance properties are associated with crosslinked network polymers, which are not currently commercially available for use in PBF AM.The defining characteristic of PBF AM as described in ISO/ASTM 52900:2021 is “thermal energy selectively fusing” specific regions of a powder bed in a layerwise manner to fabricate a 3D part [1]. When that energy is supplied via a scanning laser, the technology is termed laser- powder bed fusion (L-PBF). Due to its nature as a thermal process with solid powder-form feedstock, L-PBF is most often used to process thermoplastic polymers.

Due to its nature as a thermal process with solid powder-form feedstock, L-PBF is most often used to process thermoplastic polymers. Recent review papers that provide an overview of materials printed via L-PBF AM do not include any thermosetting polymers and often specifically describe how the heat-and-reform style manufacturing process lends itself to thermoplastic polymers [4, 2, 3, 5]. What little published literature exists for L-PBF23of thermosets can be categorized by the point in the manufacturing process relative to the printing step when the network structure is formed. The following paragraphs summarize published literature for thermoset parts made using L-PBF that form the covalent network structure either (i) pre-printing or (ii) post-printing.Covalent adaptable networks (CAN) are one way to form a network structure pre-print. CANs of poly(dimethyl siloxane) (PDMS) has been demonstrated printable via laser powder bed fusion (L-PBF) by Sun, et al. [6]. Sun, et al.

CANs of poly(dimethyl siloxane) (PDMS) has been demonstrated printable via laser powder bed fusion (L-PBF) by Sun, et al. [6]. Sun, et al. demonstrated the useful self-healing and elastomeric properties of CANs for a shoe insole application. According to their report, the PDMS CAN pre- existed L-PBF processing then was broken and reformed during printing and again in the form of a printed part to showcase its “self-healing” performance. Although CANs are covalently crosslinked networks at application temperatures, the network is broken at processing temperatures resulting in thermoplastic-like processing behavior. Such polymers do not require the additional considerations of curing while printing. Additionally, the same researchers have reported L-PBF printing a polyurethane with a dynamic covalent bond based on the Diels-Alder reaction [7]. They report dramatically decreased anisotropy in printed CANs.

They report dramatically decreased anisotropy in printed CANs. The improved z-direction tensile strength is claimed to be a result of forming dynamic crosslinks across layer boundaries. [7]The second general method for L-PBF of thermosetting polymers is forming the network post-printing. Researchers from NASA report their initial and revised printing efforts for their target aerospace grade polyimides [8, 9]. Using their method, an uncrosslinked, near-net shape “green part” is prepared using a DTM Sinterstation L-PBF machine (λ = 10.6 µm) before crosslinking the chain-ends during a 16 h post-print thermal bake. Published work by Hassan, et al. and Campbell, et al. employ similar approaches. Hassan and coauthors use a Sinterit Lisa (λ = 808 nm) to make a bismaleimide green body followed by a 61 h thermal post-process at various isothermal steps [10]. Campbell, et al.

Hassan and coauthors use a Sinterit Lisa (λ = 808 nm) to make a bismaleimide green body followed by a 61 h thermal post-process at various isothermal steps [10]. Campbell, et al. report printing a two-part epoxide plus amino sulfone green part with a Sintratec Kit (λ = 455 nm) prior to crosslinking the chain-ends during an 85 h post- print thermal bake [11].Separating the steps of macro-scale geometry formation and micro-scale polymer network formation have distinct advantages noted by authors of each prior work. Specifically, the heating environment and rheology can be tailored to the activity of interest. A mismatch between heating, flow, and curing may cause significant geometric distortion when attempting to thermally cure freestanding objects. Partially pre-reacting polymer systems provides appropriate green-part strength so that fused parts can hold shape during post-printing cure cycles.

Partially pre-reacting polymer systems provides appropriate green-part strength so that fused parts can hold shape during post-printing cure cycles. Additionally, Chung, et al.’s reported cure cycles for their polyimide involved lengthy, stepwise heating to ensure geometric stability as the crosslinked network was slowly formed in the free-standing green body [9]. Campbell, et al. discuss the importance of slow curing kinetics to avoid shape distortion [11].Note that the logical third category of forming a covalently crosslinked network during printing is absent from published literature. The present work described in this paper contributes towards closing that knowledge gap. In the present work, the authors describe translating commercial off-the-shelf (COTS) oven-cure thermosetting polymers into the local, rapid, and cyclic laser heating environment of L-PBF AM.

In the present work, the authors describe translating commercial off-the-shelf (COTS) oven-cure thermosetting polymers into the local, rapid, and cyclic laser heating environment of L-PBF AM. The discussion centers on the predicting extent of cure from in situ temperature profiles collected with an IR camera and comparing predicted values against post-print measurements. The authors conclude by synthesizing observations into guidelines and an early outlook on forming network polymers during L-PBF AM.24Materials and MethodsThermosetting Materials. One of the most challenging requirements for L-PBF AM is the spherical particle feedstock form. Most polymer L-PBF powders have an average particle diameter between 45-90 µm with a narrow overall distribution between 20-120 µm and are required to be very spherical [2].

Most polymer L-PBF powders have an average particle diameter between 45-90 µm with a narrow overall distribution between 20-120 µm and are required to be very spherical [2]. Powder paint is one current, commercial application that uses powdered thermosetting polymer precursors; therefore, the following three COTS powder paints were used in this study: Vulcan Black (TGIC-containing polyester) by Axalta, Oil Black (epoxy-based) by Axalta, and the TD 61 series Anodized Effect (TGIC-free polyester) from TIGER Drylac. These materials span three different curing reactions providing multiple crosslink pathways to evaluate in the rapidly cyclic heating environment of L-PBF.Manufacturing Methods. All specimens prepared for this work were made using a Sintratec Kit L-PBF additive manufacturing machine (λ = 455 nm), which operates in air.

All specimens prepared for this work were made using a Sintratec Kit L-PBF additive manufacturing machine (λ = 455 nm), which operates in air. The researchers determined suitable fabrication settings on the Sintratec Kit from previous experimentation to ensure complete and viable parts would be fabricated. These parameters are included in Table 1. OEM recommended isothermal oven cure times are included in SI TABLE 1 as a point of reference.Table 1.

These parameters are included in Table 1. OEM recommended isothermal oven cure times are included in SI TABLE 1 as a point of reference.Table 1. Printing parameters used for each material.MaterialChamber Temperature [°C]Powder Surface Temperature [°C]Beam Speed [mm s-1]Layer Height [µm]Hatch Spacing [µm]Number of PerimetersPerimeter offset [µm]Vulcan Black (TGIC Polyester) Oil Black (Epoxy) TD 61 (TGIC free polyester)70707050505020, 40, 6020, 40, 6020, 40, 60, 150, 235120120120808080101010111A FLIR X6901SC thermal imaging camera was used to determine the affected maximum temperature and resultant temperature decay profile from the chosen print parameter combination for each material. The camera was mounted adjacent to the Sintratec Kit to look through the FLIR Series 4 broadband crystal view port onto the powder bed. As all three tested powders are black, they were assumed to have an emissivity of 0.95.

As all three tested powders are black, they were assumed to have an emissivity of 0.95. Nine regions of interest were defined around a grid of printed cylinders. A representative image is shown in Figure 1.25Figure 1. Representative image of field of view for FLIR camera in situ temperature measurements. Each region of interest captures the maximum temperature of the circular cross- section of the printed cylinder. The grayscale image correlates to temperature readings in Celsius.Differential Scanning Calorimetry (DSC). Samples of powdered feedstock were run on a Mettler Toledo DSC 3+ unit using heating rates of 3, 5, 10, and 20 °C min-1. Test temperature ranges were varied by component material but were conducted to provide at least 25 °C of baseline measurement before and after observed phase transitions and curing reactions.

Test temperature ranges were varied by component material but were conducted to provide at least 25 °C of baseline measurement before and after observed phase transitions and curing reactions. Data analysis to integrate underneath the exothermic curing peaks and generate extent of reaction as functions of time and temperature was performed using the Mettler Toledo STARe software.Samples of printed parts were run on the same Mettler Toledo DSC 3+ unit at a heating rate of 10 °C min-1 to determine the residual cure. Complete cure (ΔHx 0) was taken to be the average of heat evolved during crosslinking of powdered feedstock from three replicate specimens at 10 °C min-1. The heat evolved over a similar temperature range in the printed parts (ΔHx,s) was 0. This ratio is “residual cure;” one minus the ratio is “extent of determined and normalized to ΔHx cure.”Isoconversional Analysis. Isoconversional analysis was performed on DSC data using both Microsoft Excel and custom Python scripts.

Isoconversional analysis was performed on DSC data using both Microsoft Excel and custom Python scripts. Data were extracted using Excel to the tabular form relating time, temperature, and integrated energy information to percent conversion. Both26linear (Excel) and non-linear (Python; Spyder IDE; SciPy curve_fit()) methods were used in combination to determine key kinetic information (i.e., activation energy and rate of reaction) of each material. Linear fitting to determine activation energy was performed via Kissinger method, shown in Equation 1 In the equation, β is the heating rate, Tp is the absolute temperature at maximum rate of reaction, E is activation energy, and R is the ideal gas constant (8.314 J mol-1 K- 1).𝑙𝑛 (𝛽 2) = 𝑇𝑝−𝐸 𝑅(1 𝑇𝑝)Equation 1Non-linear curve-fitting to determine model parameters A, m, and n was performed via Equation 2, similar to the procedure described in Reference [12].

In this equation, α is extent of reaction.ln⁡(𝛽𝑑𝛼 𝑑𝑇) + ⁡𝐸𝐴 𝑅𝑇= ln⁡(𝐴) + ⁡𝑚 ∗ ln⁡(𝛼) + ⁡𝑛 ∗ ln⁡(1 − 𝛼)Equation 2Intermediate results for linear curve fits and validation plots for non-linear curve fits areincluded as part of Supplemental Information.Adaptation of Process Parameter Bounding Theory for Thermosetting PolymersPublished research and discussion detailing process parameter value selection for thermoplastic polymers is often centered around the energy melt ratio (EMR) and EMR for degradation first put forward by Starr, et al. [13] and Vasquez, et al. [14] in their work with nylon- 12. Chatham, et al., rearranged and expanded the EMR (Equation 3 below; Equation 3 from [15]) in a manner to predict minimum and maximum laser power values for their work processing poly(phenylene sulfide) [15].

This works further adapts that equation for the case of a reactive polymer.𝑃𝑚𝑖𝑛 =ℎ × 𝑉𝑏 × 𝑙 𝑆𝐶([𝐶𝑝 × (𝑇𝑚 − 𝑇𝑏) + 𝛥𝐻𝑚] × 𝑄 × 𝛷)Equation 3In Equation 3 Pmin is the minimum power required to achieve full melting of a unit volume of powder material with the standard process parameters of hatch spacing (h), beam speed (Vb), layer height (l), and scan count (SC). The equation describes a material of density Q, powder packing fraction Φ, and specific heat Cp, being heated from the bed temperature Tb to the polymer’s melting temperature Tm. To achieve full melting, the latent heat of melting ΔHm must also be overcome.To adapt prior work that focused on the purely physical, thermoplastic process to the case of reactive polymers, we will address the following additional topics: (1) reaction activation energy and (2) temperature dependent reaction rate.

As these concerns are specific to encouraging chemical reactions during L-PBF, this new, adapted form of the EMR will be referred to as the Energy Reaction Ratio, “EXR.”27Additional Concern 1: Reaction Activation Energy. For the case of materials that undergo a reaction in their melt state, we add the activation energy for the reaction to occur EAx at temperature Tx, as shown in Equation 4. For the present work, we assume that the onset of crosslinking occurs at a temperature higher than the end of melting (i.e., Tx > Tm).𝑃𝑚𝑖𝑛 =ℎ × 𝑉𝑏 × 𝑙 𝑆𝐶([𝐶𝑝 × (𝑇𝑥 − 𝑇𝑏) + 𝛥𝐻𝑚 + 𝐸𝐴𝑥] × 𝑄 × 𝛷)Equation 4Although crosslinking is an exothermic reaction and will be expected to locally raise the temperature, it is not included in Equation 4 because this would artificially decrease the required power predicted since the printer must first overcome EA before that energy is released.Additional Concern 2: Kinetic Maxima.

When processing semi-crystalline thermoplastic polymers, the goal is to reach a fully molten state to encourage rapid coalescence. This is most 0. In reliably achieved by raising the temperature above the equilibrium melting temperature, Tm this way, the melting phenomenon follows a “higher is better” paradigm, so long as the temperature remains below degradation. Processing thermosetting polymer systems is more complex. Instead 0 threshold, there exists one or more kinetic maxima for the thermosetting system at which of the Tm the reaction proceeds at the fastest rate. Given the rapid thermal heating and decay cycles characteristic of L-PBF, we assume in this work that a kinetic maximum should be the target temperature affected by laser scanning. Affecting a temperature either below or above a kinetic maximum will limit the extent of reaction.

Affecting a temperature either below or above a kinetic maximum will limit the extent of reaction. Future work may exploit this phenomenon as a design consideration for targeting specific degrees of crosslinking or else varying the degree of crosslinking throughout the volume of the part. Equation 4 uses the generic Tx as the target temperature for crosslinking. In this work, the authors take Tx = Tx,peak to target the maximum possible extent of crosslinking.Results and DiscussionPredicted Process Parameter Range. The theoretical minimum energy density for maximum rate of reaction was determined according to the EXR (Equation 4) using the print parameter values from Table 1 and the measured material properties included in SI TABLE 1 as Supplemental Information. Figure 2 depicts the reaction rate as a function of temperature.

Figure 2 depicts the reaction rate as a function of temperature. Reaction rate is determined energetically by integrating the area under heat flow vs time data collected at various heating rates in DSC and normalizing the energy released at each timestep to the total energy evolved in the reaction. The temperature corresponding to the maximum rate of reaction is taken to be the target temperature for EXR calculation. The temperatures at maximum rate of reaction, resultant EXR material component, and theoretical maximum beam speed are presented in Table 2.28Figure 2. Rate of reaction (dα/dt) as a function of temperature at four heating rates for (a) Vulcan (TGIC polyester), (b) TDA61 (TGIC-free polyester), and (c) Oil Black (epoxy). Curves shown are the average of three replicate trials.Table 2.

Curves shown are the average of three replicate trials.Table 2. Maximum theoretical beam speeds as predicted by EXR equation targeting the temperature for measured maximum rate of reaction for three thermosetting powder paints.Materialdα/dt Peak [°C]Initial Activation Energy [kJ mol-1]EXR_material [J mm-3]Maximum Vb [mm s-1]Vulcan (TGIC Polyester) TDA61 (TGIC- free Polyester) Oil Black (epoxy)17118920521252240.2130.0970.112113248215It should be noted that the initial activation energy reported in Table 2 is used to calculate EXR as the authors assume no reaction has occurred in the one-layer thick control volume. Figure 3 depicts the change in activation energy as a function of extent of reaction. Each of the three materials studied proceeds via different reaction mechanisms yielding different degrees of variance.

Each of the three materials studied proceeds via different reaction mechanisms yielding different degrees of variance. Although initial activation energy is chosen for maximum beam speed calculation for the initially unreacted new layer, the energy requirements change over time. For a material like Vulcan (TGIC polyester), the activation energy is greatest for the unreacted state and significantly decreases as the reaction proceeds. This may result in reaching a higher temperature than desired if laser energy is not reduced as the reaction progresses. Conversely, TDA61 (TGIC-free polyester) increases the required activation energy with extent of reaction. This may result in a failure to achieve even a moderate extent of cure. The Oil Black (epoxy) material has an activation energy with the least variation, only increasing from 24 to 40 kJ mol-1 as the reaction progresses.

The Oil Black (epoxy) material has an activation energy with the least variation, only increasing from 24 to 40 kJ mol-1 as the reaction progresses. A29material with an activation energy relationship like Vulcan (TGIC polyester) may be beneficial for L-PBF application as the high initial activation energy serves as a gatekeeper preventing unwanted reaction, but the lower activation energy to continue the reaction at higher extents of conversion may enable crosslinking to occur between layers where the available energy is lower than on the surface.for likely activation energy increase in the TDA61 (TGIC-free polyester) is a kinetically network limiting morphology. At higher extents of conversion, more energy is required for molecules to diffusively move into available reaction sites. Such behavior likely also manifests as a reduction of powder coalescence rate at high extents of conversion.

Such behavior likely also manifests as a reduction of powder coalescence rate at high extents of conversion. This topic should be explored in greater detail in future work.ThecausePredicted and Measured Extent of Conversion. The temperature profiles collected at each process parameter value combination outlined in Table 1 were used to calculate extent of reaction according to Equation 5 and Equation 6. Equation 5 is calculated for values of α between 0.01 and 0.95 in steps of 0.01 using the model parameters A, m, and n determined from the fit of Equation 2. The resulting library of temperature- and conversion-dependent rate of reaction values are used in sequential evaluation of Equation 6 based on the experimentally collected T(t) data from the IR camera. In Equation 6, tf is chosen to be the time at which α = α0 + 0.01 (i.e., rate of reaction is assumed to be constant for each 0.01 conversion increment).

In Equation 6, tf is chosen to be the time at which α = α0 + 0.01 (i.e., rate of reaction is assumed to be constant for each 0.01 conversion increment). A running calculation of α after each timestep dt yields the predicted conversion vs time plot, such as the representative plots included in .Figure 3. Activation energy as a function of conversion for three studied materials. A flat-line relationship would indicate a single-step reaction mechanism for curing, which is not observed for any of the studied materials. Instead, the data suggest complex multi-step reaction pathways for each material that are also different from each other.𝑑𝛼 𝑑𝑡(𝑇,𝛼)=𝐴∗𝑒𝑥𝑝(−𝐸𝐴 𝑅𝑇)∗𝛼𝑚∗(1−𝛼)𝑛Equation 5𝑡𝑓∫𝑑𝛼 𝑑𝑡(𝑇(𝑡))𝑑𝑡⁡|𝛼= 𝛼0Equation 6𝑡030the FLIR camera only measures surface temperature of the current layer, the predicted extent of cure is also only for a single layer. Any laser effects of scanning of subsequent layers are unaccounted for in the reported extent of conversion.

Any laser effects of scanning of subsequent layers are unaccounted for in the reported extent of conversion. It should be noted that the extent of conversion levels off as the laser leaves the region of interest and the local temperature cools rapidly to temperature. This the indicates for the studied systems that laser curing only occurs during for scanning, which maintaining geometric selectivity of reaction. These are compared against extent of cure measured by DSC of parts printed at each value combination in Table 3. Looking over the data in Table 3, it is easy to notice the large variance of the reported maximum temperature. This variance cascades throughout both predicted and measured extent of cure. Although the authors have made the simplification assumptions of (i) constant distance between IR camera and every sample and (ii) a constant emissivity of 0.95, these would only effect the maximum temperature and predicted extent of conversion data columns.

The high variance in the measured extent of cure data suggests that there is a naturally high variation in extent of reaction when cured via the rapid and cyclic heating characteristic of L-PBF. Despite the high variation, the authors discuss a few key abstractions for L- PBF printing thermosetting polymers in the following paragraphs.Asreheating duringchamberis beneficialpredictionsprocessparameterFigure 4. Representative temperature profiles and resultant predicted cure profiles from Vulcan (TGIC polyester), TDA61 (TGIC-free polyester), and Oil Black (epoxy). Temperatures were collected via in situ FLIR camera at 70 °C bed temperatures and 40 mm s-1 beam speeds.31Table 3.

Temperatures were collected via in situ FLIR camera at 70 °C bed temperatures and 40 mm s-1 beam speeds.31Table 3. Measured and predicted results of three studied materials at various process parameter value combinations.MaterialVulcan (TGIC Polyester)TDA61 (TGIC free polyester)Oil Black (epoxy)EXR = 1 Target Temperature [°C]171189205Beam Speed [mm s-1]20 40 60 20 40 60 150 235 20 40 60Maximum Temperature [°C]302 ± 85 260 ± 60 178 ± 44 258 ± 51 234 ± 52 259 ± 62 153 ± 27 124 ± 7 215 ± 51 190 ± 27 194 ± 49Predicted Extent of Cure (Isoconversional Model) [%] 33 ± 52 56 ± 17 25 ± 35 94 29 32 1 ± 1 0.02 ± 0.005 2 ± 0.3 0.9 ± 0.2 0.8 ± 0.07Measured Extent of Cure (1-Hx/Hx0) [%]13 ± 7 34 ± 15 13 ± 16 1 ± 20 0.4 ± 10 4 ± 5 8 ± 10 19 ± 30 7 ± 4 9 ± 2 9 ± 1As seen in Figure 5, there is the expected inverse relationship between scan speed and maximum temperature. However, the specific relationship between temperature and beam speed is material dependent.

However, the specific relationship between temperature and beam speed is material dependent. The observed results indicate that the material with highest initial activation energy (Vulcan, TGIC polyester) has the highest maximum temperature at slower scan speeds. The trend continues with Oil Black (epoxy) being the material with lowest initial activation energy and lowest maximum temperature at the slowest scan speeds studied. Additionally, the beam speeds required to achieve the target temperatures and predicted EXR values from Table 2 and significantly slower than predicted by the EXR’s energy balance. All three studied materials required more energy (i.e., slower beam speed) than theoretically predicted to achieve the desired temperature for maximum rate of reaction. One assumption of the original EMR that is carried over into the current form of the EXR is 100% efficient conversion of laser energy into thermal energy.

One assumption of the original EMR that is carried over into the current form of the EXR is 100% efficient conversion of laser energy into thermal energy. It is likely that this is a poor assumption for the 455 nm UV diode laser used in the present work.The Vulcan (TGIC polyester) was observed to reach the highest extent of cure during L- PBF printing, as determined by post-print DSC. This is unsurprising given that the Vulcan material was observed to both reach the highest maximum temperature of the three materials at a given beam speed, and Vulcan is the only studied material to decrease activation energy as a function of32extent of reaction (See Figure 3). The exact each phenomenon however, convoluted.contributions are,of(epoxy) material is the only one where the isoconversional model underpredicts the extent of cure.

The exact each phenomenon however, convoluted.contributions are,of(epoxy) material is the only one where the isoconversional model underpredicts the extent of cure. Both the underprediction of the Oil Black material and overprediction of the TDA61 and Vulcan materials can be explained by the isoconversional verification plots shown in Figure 6. The modeled rate of reaction (solid line) for Oil Black is consistently lower than the measured rate of reaction (x’s); the opposite is true for both Vulcan (TGIC polyester) and TDA61 (TGIC-free polyester). This is especially true at 20 °C min-1, the fastest heating rate used in this study. Given that the rates observed by IR camera are on the order of 40 °C min-1, one might expect the actual deviation from the model to be even greater.The Oil BlackFigure 5. Relationship between average maximum temperature and scan speed for the three materials studied.33Figure 6.

Relationship between average maximum temperature and scan speed for the three materials studied.33Figure 6. Representative isoconversional model fit validation curves for (a) Vulcan (TGIC polyester), (b) TDA61 (TGIC-free polyester), and (c) Oil Black (epoxy). Solid lines represent measured data and dashed lines represent fitted data.Summary and Future WorkWork presented in this manuscript centers on closing the gap for fabricating a covalent networked polymer during the laser powder bed fusion additive manufacturing processing. The equations used for theoretical determination of process parameter values were expanded from their original forms that only included considerations of thermoplastic polymer process to include activation energy and acknowledge the temperature of maximum rate of reaction.

Parameter values were predicted via the new form of the Energy Melt Ratio equation, termed Energy Reaction Ratio (EXR), for three thermosetting polymer systems, each employing unique curing mechanisms, never-before fabricated via L-PBF and the resulting processability was assessed.34Temperature profiles from parameter value combinations above the predicted threshold were measured using an in situ infrared camera. Extent of reaction was predicted using an isoconversional kinetic model from the collected thermal profiles. The post-print extent of reaction measurements via DSC revealed poor agreement with the isoconversionally predicted values. Potential sources of error from measurements and model assumptions were discussed. One major concern is the degree of variation in measured temperature profiles. Future work should include an investigation in bed location dependence and scan order on maximum temperature measured.

One major concern is the degree of variation in measured temperature profiles. Future work should include an investigation in bed location dependence and scan order on maximum temperature measured. Although its size and price point make the Sintratec Kit well-suited for material development research, there are L-PBF printers (including other models from Sintratec) with more sophisticated hardware and software to ensure greater repeatability across the build platform, across layers, and across builds. Perhaps the observed temperature variation is truly a function of the interactions between the UV diode laser and studied materials. Further investigations of the same polymer systems on different L-PBF machines are required to distinguish the effects of hardware and materials from each other.There is additional opportunity to improve the isoconversional model.

As evidenced by the fit data in Figure 6, the assumptions of using only one autocatalytic and one non-autocatalytic term in the model do not accurately represent the two polyester systems studied (i.e., Vulcan and TDA61). A more robust model would include, at a minimum, a diffusive term, as the authors hypothesize this accounts for the increase in activation energy as a function of conversion for the TDA61 (TGIC-free polyester) system. Current COTS thermosetting systems designed for other manufacturing methods are often complex, multi-step reactions. We should expect future thermosetting systems designed for L-PBF processing to be equally or more complex, given the local, rapid, and layerwise cyclical thermal cycles characteristic of L-PBF.A second means of potentially improving the isoconversional predictions is to use heating rates closer to those observed in L-PBF processing when determining the kinetic triplet.

Data can be collected at such heating rates using fast calorimetry (e.g., Mettler Toledo’s Flash DSC).Future work should investigate the effect of various laser wavelengths on the extent of reaction achieved during L-PBF. It is very likely that L-PBF processing of thermosetting systems will have a higher degree of wavelength dependency than thermoplastic polymers. The additional concerns of activation energy for chemical reaction and the kinetic maximum target temperature have an increased opportunity for energy selectivity over the purely physical changes occurring during thermoplastic processing. This would be especially true when including an initiator species into the resin mixture, such as is common in vat photopolymerization AM.The future of thermosetting polymers in L-PBF hinges on the availability of powdered feedstock, which is also a concern for thermoplastic polymer L-PBF. The current work focused on COTS material available as a powder.

The current work focused on COTS material available as a powder. Most polymer processing techniques, for both thermosetting and thermoplastic polymers, do not use a powdered feedstock form. As discussed elsewhere [2], recoating requirements for L-PBF are specific and powder production is significantly more complex than, for example, converting readily available pellets into filament for FFF AM. The significant challenge of feedstock form must be addressed for L-PBF to reach its full impact.35AcknowledgementsThe authors would like to acknowledge R. Chandler Hawsey at SRNL for his assistance in DSC sample preparation and data collection. Funding for this work was provided by the NNSA Additive Coordination Team (ACT), an NA-115 driven partnership to enable additive manufacturing solutions for enterprise needs.References[1] ASTM, Additive Manufacturing - General Principles - Fundamentals and Vocabulary,ISO/ASTM, 2021.[2] C. A. Chatham, T. E. Long and C. B.

30, p. 2003062, 2020.[6] S. Sun, G. Fei, X. Wang, M. Xie, Q. Guo, D. Fu, Z. Wang, H. Wang, G. Luo and H. Xia, "Covalent adaptable networks of polydimethylsiloxane elastomer for selective lasr sintering 3D printing," Chemical Engineering Journal, vol. 412, p. 128675, 2021.[7] S. Sun, X. Gan, Z. Wang, D. Fu, W. Pu and H. Xia, "Dynamic healable polyurethane forselective laser sintering," Additive Manufacturing, vol. 33, p. 101176, 2020.[8] K. C. Chuang, T. Gornet and H. Koerner, "Challenges in Laser Sintering of Melt-Processable Thermoset Imide Resin," NASA, 2016.[9] K. C. Chuang, T. J. Gornet, K. Schneidau and H. Koerner, "Determining if laser sintering can be applied to high-temperature thermoset polyimides to enhance covalent bonding between layers through the curing of the reactive endcaps, as compared to conventional thermoplastic polymers that display poor z-directiona," Thermal Processing, 2020.[10] M. S. Hassan, K. M. M. Billah, S. E. Hall, S. Sepulveda, J. E.

28, pp. 506-516, 2019.37SUPPLEMENTAL INFORMATIONSI TABLE 1. Manufacturer recommended curing parameters and material properties for each studied material. Column headers marked with an asterisk indicate properties measured by the authors.Material OvenTemperature [°C]Vulcan Black (TGIC Polyester) Oil Black (Epoxy) TD 61 (TGIC free polyester)200130-200180-200OEM recommended Dwell Time [min] 104-13+6-25Cp* [J g-1K-1]1.71.51.4Density [g cm-3]1.631.351.2Tx,peak* [°C]171205189ΔHm* [J g-1]6.069.727.51SI FIGURE 1 Representative first heat DSC curves for three studied materials. Curves are plotted exotherm up and data were collected at a heating rate of 10 °C min-1.38SI FIGURE 2. Kissinger linear fits for determining activation energy as a function of each constant value of conversion [0.01-0.95]. Dashed line guides the eye for an example set of four datapoints from different heating rates at the same degree of conversion.39.

nufacture Of Injection Molds Using SLS.Badrinarayan, and l.W.Barlow Dept. of Chemical Engineering, The University of Texas at Austin, Austin, Texas- 78712,U.S.A.AbstractThis paper describes the use of SLS technology for the fabrication of injection mold cavities. Green shapes were made from metal - copolymer powder mixtures by SLS. The copolymer was gradually burnt out and the metal was oxidized in an air furnace. The porous oxidized metal part was subsequently infiltrated with an epoxy resin and cured. Effect of process variables in SLS, effect of oxidation cycle, dimensional changes on oxidation and epoxy infiltration of the oxidized metal part are discussed.IntroductionSelective Laser Sintering comes under the class of Solid Freeform Processes for Rapid Manufacture of parts directly from a CAD model without part specific tooling or human intervention[1,2].

This technology has been developed at the University of Texas at Austin to produce polymer parts as well as indirect metal and ceramic parts that have polymer as an intermediate binder. We have adapted SLS to manufacture indirect oxidized metal molds for injection molding applications. The traditional method for manufacturing injection molds involves precision machining of cast metals. This fabrication process is usually very complex, time intensive, expensive and also requires highly skilled labor. Rapid Prototyping using SLS offers the solution to the problem of fabricating prototypes at low costs and high speed. Time savings due to design refinements are enormous since the prototype can be fabricated in hours as compared to months and changes required can easily be made on the computer and the part rebuilt quickly.

The suitability of the material system for mold manufacture has been demonstrated in an earlier paper[3].Experimental WorkProcessing and Post-Processing: The raw material for the SLS process consists of a metal powder and a polymer powder that acts an intermediate binder. Iron powder, -325 mesh in size, was used as the metal and was obtained from Hoeganaes, Riverton, New Jersey. The polymer binder is a copolymer of poly(methylmethacrylate-co-butylmethacrylate) and has an average melt flow index of 28.9g/10min at 75 psi and 200°C[4]. The polymer powder is obtained by spray drying polymer emulsion in a Anhydro Spray Dryer. The iron powder is mixed with polymer and used in SLS for making the green shape. The polymer composition in the mixture was 40 vol % ( 9.16 wt %). SLS was used to make the green molds as well as bars for strength tests and density cups for measurement of bed density. The mold design was provided by Golden, Colorado.

SLS was used to make the green molds as well as bars for strength tests and density cups for measurement of bed density. The mold design was provided by Golden, Colorado. The surface bed temperature in SLS was maintained at 103°C to prevent powder from caking and also to keep the part from curling. The green part made by SLS was then processed in an air furnace to remove the polymer and oxidize the metal. The effect of firing conditions on oxidized metal part were studied. The effects of processing variables on part strength are described in the following section.371The porous oxidized metal preforms were infiltrated with a liquid epoxy Epoxy Infiltration: resin to increase density and strength of parts. Epoxy resins contain a reactive oxirane structure which is commonly referred to as an "epoxy" functionality. Liquid epoxy resins are converted through these reactive epoxy sites into tough, insoluble, and infusible solids.

Liquid epoxy resins are converted through these reactive epoxy sites into tough, insoluble, and infusible solids. The main issues in choosing an epoxy infiltration system are :(i) (ii) (iii)The epoxy should have low viscosity (of the order of 100 cp) at the operating temperature. The epoxy should have good mechanical properties and high heat distortion temperature. The gel time of the epoxy at the operating temperature should be large to allow sufficient the gel time is short « 1hr) the epoxy starts crosslinking time for infiltration of parts. during infiltration and this increases the viscosity and prevents infiltration. This becomes especially important when one has to infiltrate large parts.There is usually a compromise between properties. Some of the epoxy infiltration systems that were tried out were: D.E.R.

This becomes especially important when one has to infiltrate large parts.There is usually a compromise between properties. Some of the epoxy infiltration systems that were tried out were: D.E.R. 331 with DEH 24 and 331 with Ac-methyl (anhydride curing agent) from DOW Chemical Company, 1310 resin with RP1520 hardener from Ciba Giegy, and Casting resin CR 00183 from ETI Inc. Typically, the viscosity of the resin drops with temperature. The part to be infiltrated is usually heated to 60°C and the epoxy system which is also heated to approximately the same temperature is used for infiltration. Infiltration is carried out either by applying the epoxy with a paint brush or by soaking the part in epoxy till all the porosity is filled. Figure 1. shows a resin uptake curve for infiltration of a 68% porous cylindrical oxidized iron preform with D.E.R. 331 resin, Ac-methyl curing agent, and BDMA as cure accelerator at 75 °C.

Figure 1. shows a resin uptake curve for infiltration of a 68% porous cylindrical oxidized iron preform with D.E.R. 331 resin, Ac-methyl curing agent, and BDMA as cure accelerator at 75 °C. This infiltration was carried out by placing the part in a pond of the epoxy.3 .5 , , - - - - - - - - - - - - - - - - - ,3~,.j 2.51-········ .y.?;·:;.. •.. •~ .. ·.. ·.. ·1is u2-..· .);""..;11.540008000120001600020000time,secEpoxy System: DER 331 Ac methyl- BDMAInfitration Temperature75°CFigure 1. Resin Uptake Curve For Porous Oxidized Iron PreformThe maximum infiltration height was 3.25 cm. Copper and iron were used as materials for making bars to evaluate strengths.372Results and DiscussionThe first task was to select a suitable epoxy infiltration system for infiltration of porous oxidized metal preforms.

Three different epoxy resins were evaluated for infiltration using porous oxidized copper preforms fired under the same conditions. The infiltrated strengths are shown in Table 1. Even though the Ciba Giegy system shows higher strength the DOW epoxy with anhydride cure was eventually chosen because of lower viscosity and greater pot life at the operating temperature. Two different metal powders, copper and iron, were used in SLS to make bars to evaluate final strengths of the mold material. Comparison of strengths of epoxy infiltrated oxidized copper and oxidized iron bars are shown in Figure 2. All the bars were processed with identical fIring cycles and a maximum firing temperature of 300°C for 15.25 hrs. Epoxy infiltrated oxidized iron bars show much higher strengths than epoxy infiltrated oxidized copper bars and hence iron was chosen as the raw material in SLS for making the mold.InfiltrantAvg.preform density Avg Infiltrated DensityAvg. MOR (psi)(%)(%)D.E.R.

MOR (psi)(%)(%)D.E.R. 33132983841COL 131032834045CR 0018337893103Table 1. Comparison Of Epoxy InfiltrantsThe SLS of iron-copolymer mixture was carried out in a full size DTM Model 125 commercial SLS machine supplied to the University by DTM Corporation. Nitrogen was used as an inert gas during the runs. The variables in SLS manufacture of green parts are the Laser Power, the Scan spacing, the step size and step period, which determine the beam speed, and the layer thickness. These can be coupled together to form an energy term known as the scan density or Andrew number which signifies the amount of heat put in per unit area [5]. Shown below is the equation for Andrew number, An :An = P*f / (BS*SCSP)Where= An = P = f BS = SCSP =scan density, caVcm2 Power, Watts conversion factor Beam Speed, inch/sec Scan Spacing, milsThe scan density was varied by changing the beam speed between 24 and 67 in/sec while keeping the other paran1eters fixed.

The plot of strength with scan density is shown in Figure 2. The effect of density on strength is shown in Figure 4. The strength of a material usually varies as a power law function of the density as given by the equation below:=373The plot of log(a) with log(p) yields the slope "n" and intercept "A". The value of A is 7.025e-03 and n is 11.023 for the iron - copolymer SLS parts.The effect of Firing cycle on oxidized bar strength was carried out on bars made at a scan density of 3.0. The polymer decomposition and oxidation of the metal preforms were carried out in a conventional air oven. Figure 5 shows the effect of maximum firing temperature on oxidized part strength. Table 2 shows the oxidation cycle used to study the effect of temperature on oxidized part strength.Furnace Set Point CC)Time (Hrs)2005300140015001Max Temp4Table 2.

Table 2 shows the oxidation cycle used to study the effect of temperature on oxidized part strength.Furnace Set Point CC)Time (Hrs)2005300140015001Max Temp4Table 2. Oxidation Cycle To Study Effect Of Maximum Firing Temperature on PartStrength.Note: For maximum ten1peratures less than 500°C similar firing cycle is used and the ramp rates are the same.The effect of firing cycle on strength of epoxy infiltrated oxidized iron bars is shown in Table 3. We see that there is dramatic increase in infiltrated strength as we go from bars oxidized at 300°C to those oxidized at 450° C. This could be a result of higher oxidized strength of the preform fired at 450°C. The epoxy used was D.E.R. 331 with Ac-Methyl as the anhydride curing agent and BDMA as the accelerator. The parts were cured at 100°C for 2hrs and then postcured at 150°C for 6hrs. The tensile strength of the anhydride cured D.E.R. 331 epoxy is reported to be 10,000 psi[6].Max.

The build time for one mold was approximately 10hrs. The mold pieces were fired in an air oven to remove the polymer and oxidize the metal. The porous mold preform was infiltrated with D.E.R. 331 epoxy with Ac-Methyl as the curing agent and BDMA as the accelerator. The dimensional change in going from the drawing to the green part and from the green part to the final infiltrated part is shown in Figure 6. Figure 7 shows the green and infiltrated mold parts. The biggest changes are the associated with going from the drawing to the green part. The difference in dimensionality between the drawing and the green part can be removed using a beam compensating software. Most of the dimensional shrinkages on firing and post-curing are below 4%. The negative value for shrinkage, when going from the green to the infiltrated part, is the result of Itgrowth It due to oxidation of iron during the firing cycle in air.15.0~ § ..c u ........ro l::o......

I~, I, 1,1 iFigure 7. Green And Epoxy Infiltrated Oxidized Metal MoldsConclusionsSelective Laser Sintering has been used to prepare indirect metal parts that have reasonable strengths. This material system can be used to make injection molds for small scale production runs. Optimization of binder burnout and metal oxidation cycle holds the key to accurate dimensional control which is critical in production of injection molds.AcknowledgmentThe authors gratefully acknowledge support for this work by DARPA / ONR Grant NOOOI4-92-J-1394. The authors also thank N.K.Vail for his invaluable assistance in design of the mold from line drawings and in the machine operation.References1. Carl Deckard and Joseph Beaman, 15th Conference on Production Research and Technology, University of California at Berkeley, p 623 - 630, (1989).2. Harris L. Marcus, Joseph Beaman, Joel W. Barlow and David L. Bourell, JOM, 42(2), P 8-10, (1990).3. James R. Tobin, B. Badrinarayan, 1. W. Barlow, 1.

EE FORM FABRICATION OF METALLIC COMPONENTS USING LASER ENGINEERED NET SHAPING (LENS™)M. L. Griffith, D. M. Keicher, C. L. Atwood, J. A. Romero, 1. E. Smugeresky*, L. D. Harwell, D. L. GreeneSandia National Laboratories Albuquerque, NM 87185, *Livermore, CA 94551INTRODUCTIONSolid free form fabrication is one of the fastest growing automated manufacturingtechnologies that has significantly impacted the length of time between initial concept and actual part fabrication 1 2. Starting with CAD renditions of new components, several techniques such as stereolithography3 and selective laser sintering4 are being used to fabricate highly accurate complex three-dimensional concept models using polymeric materials.

Coupled with investment casting techniques, sacrificial polymeric objects are used to minimize costs and time to fabricate tooling used to make complex metal castings5This paper will describe recent developments in a new technology, known as LENSTM (LaserEngineered Net Shaping)6 7 8 9, to fabricate metal components directly from CAD solid models and thus further reduce the lead times for metal part fabrication. In a manner analogous to stereolithography or selective sintering, the LENSTM process builds metal parts line by line and layer by layer. Metal particles are injected into a laser beam, where they are melted and deposited onto a substrate as a miniature weld pool.

Metal particles are injected into a laser beam, where they are melted and deposited onto a substrate as a miniature weld pool. The trace of the laser beam on the substrate is driven by the definition of CAD models until the desired net-shaped densified metal component is produced.EXPERIMENTALThe system consists of aNd:YAG laser, a controlled atmosphere glovebox, a 3-axis computer controlled positioning system, and a powder feed unit. The positioning stages are mounted inside a controlled atmosphere glove box, backfilled with argon, operating at a nominal oxygen level of 2-3 parts per million. The beam is brought into the glovebox through a window mounted on the top of the glovebox and directed to the deposition region using a six inch focal length plano-convex lens.

The beam is brought into the glovebox through a window mounted on the top of the glovebox and directed to the deposition region using a six inch focal length plano-convex lens. The powder delivery nozzle is designed to inject the powder stream directly into the focused laser beam and the lens and powder nozzle move as an integral unit.125A CAD solid model is sliced into a sequence of layers, and translated into a series of tool path patterns to build each layer. Each layer is fabricated by first generating an outline of the key component features and then filling the cross-section using a rastering technique. This file is used to drive the laser system to produce the desired component one layer at a time, starting from the bottom of the part. A schematical representation of the LENSTM fabrication process is shown in Figure 1. A solid substrate is used as a base for building the LENSTM object.

A schematical representation of the LENSTM fabrication process is shown in Figure 1. A solid substrate is used as a base for building the LENSTM object. The laser beam is focused onto the substrate to create a weld pool in which powder particles are simultaneously injected to build up each layer. The substrate is moved beneath the laser beam to deposit a thin cross section, thereby creating the desired geometry for each layer. After deposition of each layer, the powder delivery nozzle and focusing lens assembly is incremented in the positive Z direction, building a three dimensional component layer additively.

After deposition of each layer, the powder delivery nozzle and focusing lens assembly is incremented in the positive Z direction, building a three dimensional component layer additively. To insure that a uniform deposition was achieved for each layer independent of direction, a specialized powder delivery nozzle and powder feeder have been developed9 .Z·axis Positioning of Focusing Lens and Powder Delivery NozzleLaser Beam1.41.2Powder DeliveryNozzleE .s :c '" .;;; :r: c ,g iii 0 Co"c0.80.80.40.2•••0"X-V Positioning Stages10152025304 ~.Power DensityfTraverse Velocity (J/cn1)Figure 1: Schematic of the LENS process.Figure 2: Build height as a function of the volumetric exposure.RESULTS A. LaserlPowder Interactions (i) Process Parameters Our initial research was to demonstrate the role of the processing parameters for thin walledgeometries.

LaserlPowder Interactions (i) Process Parameters Our initial research was to demonstrate the role of the processing parameters for thin walledgeometries. Statistically designed experiments were done using Inconel 625 to identify significant process variables and to begin to understand the deposition process variables considered were: component velocity, laser irradiance, Z-axis increment, and powder volumetric flow rate. The response variables used in evaluating the experimental results were: the material build-up height, the melt depth into the previous layer, and the ratio of these two variables. The tests were performed by depositing ten layers of metal in a single wall for each of the experimental conditions, where the material was deposited in only one direction of travel. Metallographic analysis was used to measure the response variables and to evaluate the experiment results.

Metallographic analysis was used to measure the response variables and to evaluate the experiment results. Analysis of this data for the Ni-based super alloy 625 has shown that there is a linear relationship between the layer build-up height and total volumetric exposure. The6The process126volumetric exposure is the ratio of the laser irradiance to the component travel speed. A graph of these experimental results is shown in Figure 2.Complete melting of the powder occurred for all tests. In addition, textured growth of thedeposited material occurred across the deposition layer boundary in nearly all cases for thin walled parts. The build up height was measured from the original surface to the top of the deposited structure. Similarly, the melt depth was taken to be the depth of the dissolution region. Previous work6 intergranular melting which has occurred was only a fraction of the substrate grain size.

Similarly, the melt depth was taken to be the depth of the dissolution region. Previous work6 intergranular melting which has occurred was only a fraction of the substrate grain size. This suggests that for the conditions used in this testing the heat affected zone is relatively small.,8 shows there was little intergranular melting in the substrate region. TheFor each new material processed by the LENS technique, test matrices are set up to determine the correct parameters to build near net parts. This has been done for a variety of materials: stainless steels, nickel-based alloys, HI3 tool steel, and some tungsten.

This has been done for a variety of materials: stainless steels, nickel-based alloys, HI3 tool steel, and some tungsten. After thin walled parts are built, solid block test matrices are performed to understand the inter- and intra layer parameters needed to build fully dense parts.(ii) Diagnostics Several diagnostic techniques are used to monitor and to understand the LENS process.These include: entering the laser beam10 magnification, high speed digital imaging for process understanding6 imaging.laser Doppler velocimetry for powder flow and relative density of the powder, time resolved infrared imaging for thermal characteristics 11, high, and standard videoHigh speed, high magnification imaging allowed the LENSTM process to be effectively slowed down to visualize the molten metal/powder interaction region. It appears that particles do not become molten until they are, in fact, actually injected into the melted metal puddle in the deposition region.

It appears that particles do not become molten until they are, in fact, actually injected into the melted metal puddle in the deposition region. Two powder sizes were used to study the process and weld pool behavior. For the smaller particle size (-325 mesh), the melt puddle appear to be stable and well behaved. For the larger particle size distribution (-80 to +325 mesh), the molten puddle was very energetic and unstable. For the larger powder size distribution, the particle size was a significant fraction of the deposition region width. Directing the larger particles into the molten deposition region causes a larger displacement of the liquid metal thus adding more energy to the oscillations of the melt pooL Further studies are required to draw more quantitative conclusions for the effects of particle size on the powder deposition process.B. Three-dimensional componentsAfter determining the basic LENSTM parameters for a material, a hollow geometry is typicallyfabricated.

Three-dimensional componentsAfter determining the basic LENSTM parameters for a material, a hollow geometry is typicallyfabricated. Figure 3 is a picture of an HI3 tool steel part. The tallest geometry we can build in the LENSTM platform is 6 inches. For this geometry and material, the dimensional variance along Z is only 0.002 inches for the wingspan section of the thunderbird. The surface finish on extruded shapes is typically 250 Jlinch.From our understanding of the LENSTM parameters, solid geometries are fabricated. With asolid geometry, understanding the hatch spacing (line by line spacing) is critical for building127100% dense components. Figure 4 shows a housing built out of 316 stainless steel (SS316). This part is very accurate in X and Y dimensions but the build height needs better control. The error in X and Y is less than ±0.005"; however, Z can vary by as much as ±0.015" where the substrate extracts the thermal energy from the first few layers.

The error in X and Y is less than ±0.005"; however, Z can vary by as much as ±0.015" where the substrate extracts the thermal energy from the first few layers. By understanding the thermal behavior along with the LENSTM parameters, we should be able to control the Z height.Figure 3: Thin-walled geometry fabricated from H13 tool steel.Figure 4: Solid geometry fabricated from 316 stainless steel.Currently this process is being developed as a free-form fabrication process in which no support structures have been needed. Preliminary results from angle build studies suggest that the maximum angle which can be achieved in a single width deposition is approximately 30° and about 15° for solid parts.

Preliminary results from angle build studies suggest that the maximum angle which can be achieved in a single width deposition is approximately 30° and about 15° for solid parts. A horizontal rotation axis has been added to the LENS system and will expand the geometries which can be built.Although visual analysis of sample cross-sections exhibited no obvious signs of porosity,helium pycnometry, Archimedes' method, and ultrasonic imaging were used to quantify the density for LENS processed materials. Both the pycnometry and Archimedes method showed the LENS processed materials to be fully dense. Ultrasonic imaging showed that the SS316 processed material had a few microvoids on the order of 1-5 J.lm in the interior of the solid. Electron microprobe analysis indicated that there was no apparent difference in composition between the deposited material and the original substrate (which is same material).c.

Electron microprobe analysis indicated that there was no apparent difference in composition between the deposited material and the original substrate (which is same material).c. AccuracyPhotographs of tolerance test parts used to measure the dimensional accuracy of the LENS™ process are shown in Figures 5 and 6. In building these components, it was determined that the dimensions in the X-Y plane could be maint.ained to less than ±0.002 inches (0.02 mm). The dimension in the Z or growth direction could only be maintained within ±0.015 inches (0.4 mm). The angle in the pyramid of Figure 6 is maintain within ±0.015 inches (0.4 mm). These results are extremely promising and further work will allow improvements in the dimensional accuracy to be achieved. The surface finish appears to be a strong function of the powder particle size with the smaller powder particles giving a better surface finish.

The surface finish appears to be a strong function of the powder particle size with the smaller powder particles giving a better surface finish. Although the surface finish of the "as128built" component is somewhat rough, a small amount of finishing will produce an accurate, high polished surface.Figure 5: Accuracy test part with step geometry and cylindrical holes.Figure 6: Square-circle-pyramid accuracy test part.D. Mechanical TestingResults from preliminary tensile testing of the deposited 316 stainless steel are given in Table 1. Eight bars have been tested to date, six with the layers perpendicular to the tensile pull direction, and two samples with the layers parallel to the tensile pull direction. A third set of data is given for conventionally processed, annealed 316 stainless steel12 strength properties of the fabricated 316 stainless steel bars significantly exceeds that for the reported value of the annealed material.

The elongation for the LENS fabricated have similar ductility where an elongation of 50% (in 1 inch) is achieved. Particle size did not effect the strength properties; but some of the early tensile samples were fabricated with a glove box oxygen content around 20-50 ppm, which lead to premature reduction in ductility. With low oxygen content « 2 ppm), the ductility values were always greater than 50%.In all cases, theTable 1: Mechanical tensile test results for LENS fabricated SS316 bars. Plane orientation represents the layer build style with respect to the tensile pull direction. Yield Strength (ksi) 65Plane Orientation (w.rt. tensile direction) perpendicular (-325 mesh) perpendicular (+100/-325 mesh) parallel (-325 mesh) annealed bar lUltimate Tensile Strength (ksi) 115Elongation% (in 2.54 cm) 36-6665115518611733£358550Metallographic cross-sections of volumetric LENS fabricated parts exhibit no textured grain growth across the deposition layers.

Moreover, Poisson's ratio is isotropic, and the tensile results do not show any preferred properties for specimens made with the layers perpendicular or parallel to the pull direction.129Preliminary work with HI3 tool steel shows that it is readily possible to make componentswith a hardness value of 59.3 (Rockwell C).SENSORSAlthough this process has proven to be very robust, a significant amount of work remains to develop the LENSTM process for operation in a manufacturing environment. The current system operates as an open-loop system, where it depends solely on the reliability of the laser and other system components to reproduce a given result. Preliminary results from current experimentation suggests that improvements can be made in process reliability by implementation of sensors for monitoring the process and providing a response signal for closed loop process control.Research in the area of sensors to monitor the LENS process and for closed loop control isongoing.

Sensors being developed include: powder mass flow for accurate flow conditions I , thermal monitoring of the weld pool and the workpiece to obtain uniformity in fabrication and resulting material's properties I I, pyrometry to monitor characteristics of the weld pool for feedback control I I, and in-situ, z-height determination of the workpiece in relation to the nozzle to obtain accurate parts.CONCLUSIONSThe feasibility of fabricating fully dense, solid metallic components directly from a CAD solid model has been demonstrated. The material properties obtained using alloys such as the Inconel 625 and 316 stainless steel are comparable to a conventionally processed wrought material, in some cases the material properties obtained in the LENSTM fabricated structure far exceed those for annealed materials.

Dimensional studies have shown that very precise tolerances can be achieve in the horizontal build plane and the data generate from these studies has suggested ways to control and improve dimensional accuracy in the vertical fabrication direction. Further work is underway to continue to improve the material surface finish and modified fabrication techniques are being explored to overcome angular limitations imposed by the current fabrication approach.ACKNOWLEDGMENTWork supported by the U. S. Department of Energy under contract DE-AC04-94AL85000.REFERENCES1. M. Bums, Automated Fabrication: Improving Productivity in Manufacturing, Prentice Hall, Englewood, NJ, (1993).2. T. Wohlers, "State of the Industry", Proceedings of the Rapid Prototyping & Manufacturing '96 Conference, published by the Society of Manufacturing Engineers, Dearborn, MI, (1996).13003. P.

T. Wohlers, "State of the Industry", Proceedings of the Rapid Prototyping & Manufacturing '96 Conference, published by the Society of Manufacturing Engineers, Dearborn, MI, (1996).13003. P. Jacobs, Rapid Prototyping and Manufacturing: Fundamentals of Stereolithography, published by the Society of Manufacturing Engineers, Dearborn, MI, (1992).4. K. McAlea, R. Booth, P. Forderhause, U. Lakshminarayan, "Materials for Selective Laser Sintering Processing", 27th International SAMPE Technical conference, Vol. 27, Diversity into the Next Century, proceedings ofSAMPE '95, Albuquerque, NM, October 12-14, 1995, p. 949.5. M. D. Baldwin, C. L. Atwood, and M. C. Maguire, "Integration of Rapid Prototyping into Investment Casting", Proceedings of the Rapid Prototyping & Manufacturing '95 Conference, published by the Society of Manufacturing Engineers, Dearborn, MI, (1995).6. D. M. Keicher, J. L. Jellison, L. P. Schanwald, J. A. Romero and D. H.

eeform Bioprinting of Liver Encapsulated in Alginate Hydrogels Tissue Constructs for Pharmacokinetic Study Chang, R., Starly*, B., Sun, W. Dept. of Mechanical Engineering, Drexel University, Philadelphia, PA 19104 Culbertson, C., Kansas State University, Manhattan, KS 66506 Holtorf, H., Gonda, S., NASA JSC, Houston, Texas 77058Abstract An in vitro model that can be realistically and inexpensively used to predict human response to various drug administration and toxic chemical exposure is needed. By fabricating a microscale 3D physiological tissue construct consisting of an array of channels and tissue-embedded chambers, one can selectively develop various biomimicking mammalian tissues for a number of pharmaceutical applications, for example, experimental pharmaceutical screening for drug efficacy and toxicity along with apprehending the disposition and metabolic profile of a candidate drug.

This paper addresses issues relating to the development and implementation of a bioprinting process for freeform fabrication of a 3D cell-encapsulated hydrogel-based tissue construct, the direct integration onto a microfluidic device for pharmacokinetic study, and the underlying engineering science for the fabrication of a 3D microscale tissue chamber as well as its application in pharmacokinetic study. To this end, a prototype 3D microfluidic tissue chamber embedded with liver cells encapsulated within a hydrogel matrix construct is bioprinted as a physiological in vitro model for pharmacokinetic study. The developed fabrication processes are further validated and parameters optimized by assessing cell viability and liver cell phenotype, in which metabolic and synthetic liver functions are quantitated.1. IntroductionTissue engineering approaches exploit living cells in a variety of ways towards the goal of restoring, maintaining, or enhancing tissues and organs [1,2].

IntroductionTissue engineering approaches exploit living cells in a variety of ways towards the goal of restoring, maintaining, or enhancing tissues and organs [1,2]. In order to engineer biological tissues in vitro, cultured cells are coaxed to grow on bioactive resorbable scaffolds, i.e. temporary synthetic extracellular matrices that provide the biological, chemical, and mechanical cues to guide the cell’s eventual differentiation and assembly into three-dimensional (3D) tissues [3]. While regeneration of different tissues and organs are currently under heavy investigation and development, new applications of tissue engineering in designing in vitro physiological models to study disease pathogenesis and for pharmacokinetic study are also extremely promising [4]. One possible application of in vitro physiological models established from biology-based engineering analysis is in the area of pharmaceutical drug screening and new drug discovery and development.

One possible application of in vitro physiological models established from biology-based engineering analysis is in the area of pharmaceutical drug screening and new drug discovery and development. In vitro cell culture models with human liver cells have already shown great potential in predicting studies on drug toxicity and metabolism in the pharmaceutical industry [5,6]. For example, Zeilinger et al developed a bioreactor culture model that permits the 3D coculture of liver cells under continuous medium perfusion with decentralized mass exchange and integral oxygenation [7]. Powers and Domansky designed a microfabricated array bioreactor for perfused 3D liver culture at reported perfusate flow rates and fluid shear stresess at or below the physiological shear range in vivo.

These primary rat liver cells cultured for two weeks in the channels rearranged themselves to form tissue-like structures.Current address: School of Industrial Engineering, University of Oklahoma, Norman, OK 73019138Shuler et al has conceived a microscale in vitro system to serve as a human surrogate for drug analysis which adheres stringently to a physiologically-based pharmokinetic (PBPK) model in which disparate organs are serially connected by channels to comprise a fluid circuit. This methodology is then used to mechanistically simulate and predict the fate of drugs or other substances in vivo.

This methodology is then used to mechanistically simulate and predict the fate of drugs or other substances in vivo. Therefore, one can conceive and fabricate a stamp-sized animal-on-a-chip microfluidic device which, with great fidelity, replicates an experimental drug’s being broken down by the metabolizing liver, absorbed by the intestines, and held onto by fat [8-10].The present work explores the development and study of in vitro three-dimensional Microfluidic Microanalytical Microorgan Device (3MD) for simulation of physiological human response to drug administrations and toxic chemical exposure. The applied solid freeform fabrication technology is a viable bioprinting freeform fabrication process for layer-by-layer extrusion of 3D cell-encapsulated hydrogel-based tissue constructs.

The applied solid freeform fabrication technology is a viable bioprinting freeform fabrication process for layer-by-layer extrusion of 3D cell-encapsulated hydrogel-based tissue constructs. By fabricating a 3D in vitro tissue analog consisting of an array of channels with tissue-embedded chambers, one can selectively biomimic different mammalian tissues for a multitude of applications, foremost among them liver tissue for clinical pharmaceutical screening of drug efficacy and toxicity. The research conducted is aimed at the achievement of high-throughput reproducible fabrication of bioprinted tissue constructs and 3D organ chambers, maintenance of structural integrity and the integration with the microfluic platform, and enhancement of cell viability and cell phenotype retention.

More specifically, the objectives are to develop a viable bioprinting process for fabrication of reproducible 3D cell-encapsulated alginate-based tissue constructs; to fabricate 3D organ chambers, i.e., bioprinted liver tissue constructs with surface-treated silicon microfluidic devices; and to study the effect of the process on cell viability and cell-specific function retention.2. System Configuration for 3D Tissue Constructs Within Microfluidic Device2.1 OverviewThe bioprinting process is integrated with a microfluidic device to fabricate 3D tissue/organ constructs/chambers, as opposed to producing 2D cell monolayers (Figure 1). Biological studies reveal the instability cellular phenotype and reduced tissue-specific gene expression with conventional monolayer in vitro culture techniques [11-13].

Biological studies reveal the instability cellular phenotype and reduced tissue-specific gene expression with conventional monolayer in vitro culture techniques [11-13]. A three-dimensional tissue model will, in contrast, foster improved retention of hepatocyte-specific function.Figure 1: Overview of Bioprinting Tissue-on-a-Chip Approach139Furthermore, direct cell deposition of encapsulated cells offers tighter control over the spatial distribution of cells and doesn’t rely upon cell migration to populate the scaffold, allowing one to assembly high cell density or co-culture multiple cell types within a 3-dimensional construct. This can create tissue structures that more closely resemble their in vivo state. Cell-cell communication either from direct contact or paracrine signaling is important for proper cellular behavior, differentiation, and proliferation, along with the concomitant extracellular matrix produced by the neighboring cells.

Furthermore, optimization of process parameters (e.g. nozzle pressure, motion arm velocity, nozzle tip size etc.) and material parameters (e.g. biopolymer viscosity, crosslinking agent concentrations, etc.) have been done to achieve high-fidelity 3D structures and seamless integration onto a microfluidic tissue micro-organ chambers.2.2 Modifications of System for Tissue-on-a-chip ApplicationFirst, mouse hepatocytes are encapsulated within alginate hydrogels and bioprinted through the system into a desired pattern within the microfluidic circuits. The other component to the microfluidic microanalytical microorgan device (3DM) system is the micofluidic device with indented chambers. The feasibility of using a standard soft-lithography technique is explored to fabricate microscale in vitro device with microchannels and multi-chambers to house the bioprinted tissue. The preferred material for fabrication of the microchip for housing the embedded tissue is PDMS.

The preferred material for fabrication of the microchip for housing the embedded tissue is PDMS. PDMS elastomer soft lithography is combined with the micro- molding techniques to fabricate 3D micro-fluidic chambers. The advantages of these fabricated chambers are the ease of bonding, optical properties, and permeability to gases for biological application.Figure 2: Setup of Bioprinted Tissue-on-a-chip for Medium/Drug Flow StudyThe next step is to bioprint the the cell-encapsulated alginate constructs within a PDMS chambers. The PDMS chambers contain a 10mm x 10mm x 0.6mm (Width x Length x Height) indentation. In order to integrate the hydrophilic alginate constructs onto the PDMS substrate, UV treatment was applied to the PDMS chips to oxidize the surface methyl groups to form silanol groups along with roughening the surface. Therefore, a hydrophilic surface with good wettability and improved traction is attained.

Therefore, a hydrophilic surface with good wettability and improved traction is attained. Next, the cell-encapsulated alginate construct is directly bioprinted within the tissue chamber indentation on the PDMS microchip, thus forming140an integrated 3-D tissue chamber unit. Figure 2 shows a setup of applying such tissue-on-a-chip unit for medium/drug flow study.A schematic of medium/drug flow circulation in the mini-bioreactor is shown in Figure 3. To demonstrate effective drug metabolism in the liver chamber, a non-fluorescent prodrug is fed into the system through the inlet port, metabolized by the liver chamber, and then produces an effluent fluorescent metabolite for analysis collected at the outlet port. Results of such analysis can be used to understand the relative pharmacokinetic efficiency as well as relevancy of the tissue chambers design for human application.Figure 3: Schematic of Medium/Drug Flow Circuit through tissue chambers3.

Bioprinting on a Chip Pattern3.1 Overview of existing bioprinting systemA proprietarily developed Drexel multinozzle bioprinting system is capable of depositing heterogeneous materials, cell types, and biological factors in a controlled and reproducible manner [14,15]. This bioprinting process employs solid freeform fabrication (SFF) techniques in conjunction with computer-aided modeling of heterogeneous structures to build biopolymer- based 3D cell-embedded tissue constructs. The deposition system utilizes microvalve nozzle systems that can deposit a wide range of solutions with a wide range of material and biological properties. The system configuration allows a pre-specified computer-aided design (CAD) to be converted into a layered process toolpath to extrude biopolymeric materials, cells, and other biological factors through various nozzle systems.

An overview of the bioprinting system configuration is shown in Figure 4.3.2 Fluid dynamical modeling of the bioprinting systemThe bioactive fabrication process described herein involves extruding live cells admixed homogeneously with sodium alginate through the pneumatic system. In the deposition, the cells are exposed to the shear mechanical perturbations within the cell suspension that could potentially permanently kill or temporarily damage cells, thereby altering functionality (e.g., drug metabolism). Therefore, an expression for the shear stress would be helpful in quantifying141the amount of shear that cells may experience as they traverse the pneumatic-driven bioprinting nozzle system.Figure 4: Overview of System Configuration of 3D Bioprinting SystemFundamental equations from fluid dynamics can be integrated into an analytical model of the mechanical forces on cells and biomaterial flowing through a nozzle tip.

(cid:74)is the shear rate, R is the nozzle tip radius, constant K is thewhere (cid:536) is the viscosity, consistency index, and the constant n is the power law index that is equal to unity for Newtonian liquids and is less than one for non-Newtonian liquids [16].3.3 Bioprinted 3D Tissue Construct Structural FormabilityA micro-scale liver tissue analog has been designed and fabricated for studying drug metabolism and its pharmacokinetic effects on secondary organ function via direct deposition of a three dimensional heterogeneous cell-seeded hydrogel-based matrix. By integrating the bioprinting system with a CAD environment, notable feasibility and reproducibility of 3D structures has been realized within micron-order dimensional specifications and hence can be used for parametric studies (Figure 5).Figure 5: Bioprinted Tissue Construct Structural Formability4.

Preliminary Study of Cell Viability and Hepatocyte-Specific Function4.1 Preliminary Study of Cell ViabilityPreliminary cell viability assays were conducted to study the effect of the bioprinting process on cell survival. The aim of this study was to evaluate the cell viability ratio at Day 0 of the bioprinting system. The experimental setup for this study (Figure 6) involved 2 test samples and one control. Test 1 was the bioprinting of alginate and hepatocytes with cell medium serving as143the cross-linking agent. Test 2 was the bioprinting of alginate and hepatocytes with a CaCl2 polymerizing cross-linker. Both the process and material parameters for Test 1 and Test 2 were held constant. The control implemented was manual pipetting of alginate and hepatocytes with cell medium as the cross-linking agent.

The control implemented was manual pipetting of alginate and hepatocytes with cell medium as the cross-linking agent. The same cell density (175,000 cells/mL) was used in both tests and the control.Figure 6: Post-Processing Hepatocyte Cell ViabilityFor this setup, qualitative and quantitative assays were carried out with Live/Dead cell assay and Alamar Blue cell assay respectively. Each data point represented samples taken at Day 0 shortly after bioprinting and subsequent cross-linking. The rationale for this setup with assay was to compare and isolate the effects of the bioprinting process and biomaterial cross-linking agent on the cell viability.

The rationale for this setup with assay was to compare and isolate the effects of the bioprinting process and biomaterial cross-linking agent on the cell viability. Preliminary testing demonstrated good initial cell viability of post-assembly bioprinted encapsulated hepatocytes under biofriendly conditions with Live/Dead cell assays and Alamar Blue staining (Figure 8).Hepatocyte Cell Viability as a Function of Process Parameters85%83%80%79%75%70%70%Medium Only Printing+Medium Printing + CaCl265%60%Process Parameters(a) Live-Dead Cell Assay(b) Alamar Blue Quantitative Cell Viability AssayFigure 7: Post-Processing Hepatocyte Cell ViabilityFigure 7a is a qualitative assessment at Day 0 of the cell viability with metabolically live cells denoted in green and metabolically inactive dead cells in red. Figure 7b shows the cell viability ratios at Day 0 obtained from quantitative fluorescent measurements of the two test setups and control.

Figure 7b shows the cell viability ratios at Day 0 obtained from quantitative fluorescent measurements of the two test setups and control. The control with no bioprinting and cell medium cross-linker showed 83% Test 1 bioprinting alginate/hepatocytes with cell medium cross-linker cell viability.144demonstrated a 79% cell viability. Test 2 bioprinting alginate/hepatocytes with CaCl2 cross- linker gave a 70% cell viability. A comparison of Test 1 and control indicated a slight drop in cell viability may be attributed to the bioprinting process. Further, a comparison of Test 2 and Test 1 indicated that a larger drop in cell viability may be attributed to the substitution of a stronger cross-linker (CaCl2) versus one with trace amounts of cross-linking ions (cell medium).

Overall, the cell viability tests showed that hepatocytes were able to survive through bioprinting process with a range of 70% to 79% viability ratio depending on choice of chemical cross- linking agent.4.2 Preliminary Result of Urea Synthesis for Hepatocyte FunctionA 3-day preliminary study was performed to investigate differential hepatocyte-specific function of urea synthesis in bioprinted 3D alginate encapsulated hepatocytes and 2D hepatocyte cell monolayer. The aim of this study was to evaluate the implications of a bioprinted 3D structure on a measure of hepatocyte-specific function, namely urea synthesis. The experimental setup for this study involved one experimental test sample and one control. The experimental test sample was the bioprinting of a 3D hepatocyte/alginate tissue construct. The control used was a 2D hepatocellular monolayer with the same cell density (175,000 cells/mL) as the test sample.

The control used was a 2D hepatocellular monolayer with the same cell density (175,000 cells/mL) as the test sample. In Figure 8, urea synthesis is quantitatively detected with a cell-based colorimetric assay, allowing comparison of time-dependent phenotype for bioprinted cell-encapsulated hepatocytes with controls cultured as a hepatocellular monolayer. Cell medium was replenished for both the test sample and control once per day. Each data point was an average of 5 samples taken at Days 1, 2, and 3 after bioprinting.Figure 8: Results of 3-day Urea Synthesis Study of Bioprinted Alginate Tissue Constructs of Microencapsulated HepatocytesThis data suggests that a hepatocytes encapsulated in a 3D alginate/hepatocyte tissue contruct synthesizes a higher amount of urea than the same number of hepatocytes cultured as a 2D hepatocyte cell monolayer.

This difference wass most striking on Day 1, where the 3D tissue construct showed a marked urea concentration of 1.90mg/mL compared to a 2D hepatocyte cell145monolayer urea concentration of 0.65mg/mL. While the marked dropoff in urea synthesis at Day 3 was typical of in vitro culture, the 3D structure conferred an ability for increased urea production relative to the 2D monolayer case. Therefore, this demonstrated that the bioprinting process for microencapsulation of hepatocytes in 3D alginate tissue constructs was compatible with prolonged maintenance of hepatotocellular specific function.5. Discussion and ConclusionsOver the last four years, the development of integrated devices that combine cell culture and microfabrication makes the possibility of commercial applications to pharmaceutical evaluation a real possibility.

The research undertaken herein rests on the premise and conviction that an in vitro model that can realistically and inexpensively test the response of humans and animals to various chemicals is needed. Such an in vitro system may serve to improve our ability to predict animal and human response to drug and chemical exposure. Process development of a proprietary bioprinting system to fabricate tissue construct for direct integration on a microfluidic chip will enable a novel creation of a tissue analog for pharmacokinetic study. Bioprinting tissue constructs with 3D architecture and microscale design features has been demonstrated for both feasibility and reproducibility. Cell viability tests showed that hepatocytes were able to survive through bioprinting process with a range of 70% to 79% viability ratio. Furthermore, biological assays for liver-specific function showed a conferred advantage of a 3D tissue construct over 2D cell monolayer for maintenance of cell-specific function.

Furthermore, biological assays for liver-specific function showed a conferred advantage of a 3D tissue construct over 2D cell monolayer for maintenance of cell-specific function. Ongoing research and study are being done to enhance feasibility of bioprinted 3-dimensional structural formation with reproducibility, integration of bioprinted tissue onto a microfluidic device with no leakage, cell viability, and maintenance of cell-specific function. By taking advantage of the unique bioprinting system capabilities, advanced research topics can be pursued in tissue engineering and its downstream applications, specifically pharmacokinetic study for drug screening and development. The potential impact of the tissue-engineered model physiological surrogate system developed herein will hinge on a deepened appreciation of the engineering and scientific basis for the fabrication as well as the relation with subsequent tissue pharmacokinetic behaviorReferences1. Langer R.

eeform Extrusion Fabrication of Titanium Fiber Reinforced Bioactive Glass ScaffoldsAlbin Thomas1, Krishna C.R. Kolan1, Ming C. Leu1 and Gregory E. Hilmas21Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409 2Department of Material Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractAlthough implants made with bioactive glass have shown promising results for bone repair, their application in repairing load-bearing long bones is limited due to their low fracture toughness and fairly fast degradation response in vivo. In this paper, we describe our investigation of freeform extrusion fabrication of silicate based 13-93 bioactive glass scaffolds reinforced with titanium fibers. A composite paste was prepared with 13-93 bioactive glass filled with titanium fibers (~16 µm in diameter and aspect ratio of ~250) having a volume fraction of 0.4 vol. %.

A composite paste was prepared with 13-93 bioactive glass filled with titanium fibers (~16 µm in diameter and aspect ratio of ~250) having a volume fraction of 0.4 vol. %. This paste was then extruded to fabricate scaffolds with an extrudate diameter of about ~0.8 mm. The sintered scaffolds, with and without titanium fibers, had measured pore sizes ranging from 400 to 800 µm and a porosity of ~50%. Scaffolds produced with 0.4 vol. % titanium fibers were measured to have a fracture toughness of ~0.8 MPa•m1/2 and a flexural strength of ~15 MPa. Bioactive glass scaffolds without titanium fibers had a toughness of ~ 0.5 MPa•m1/2 and strength of ~10 MPa. The addition of titanium fibers increased the fracture toughness of the scaffolds by ~70% and flexural strength by ~40%. The scaffolds’ biocompatibility and their degradation in mechanical properties, in vitro were assessed by immersing the scaffolds in a simulated body fluid over a period of one to four weeks.1.

The scaffolds’ biocompatibility and their degradation in mechanical properties, in vitro were assessed by immersing the scaffolds in a simulated body fluid over a period of one to four weeks.1. IntroductionBioactive glasses have been widely investigated for bone repair applications, because of their inherent capability to enhance new bone formation and the ability to bond to surrounding tissues [1]. Out of a variety of bioactive glasses, 13-93 bioglass is a silicate based glass (53 SiO2, 20 CaO, 4 P2O5, 6 Na2O, 12 K2O, and 5 MgO, with chemical composition in weight. %) which has been thoroughly researched in the past and FDA approved for its proven chemical durability, bioactivity, bone cell proliferation and differentiation and more importantly its ability to be fabricated into scaffolds.

Fu et al.[2], Liu et al.[3], Kolan et al.[4], Doiphode et al.[5], and Rahaman et al.[6] studied the use of 13-93 bioglass in fabricating scaffolds for bone repair using a variety of fabrication techniques including slip casting, polymer foam replication, selective laser sintering, freeze extrusion fabrication, and robocasting, respectively . The investigation into mechanical and biological properties of the fabricated scaffolds has shown that 13-93 bioglass has favorable properties in repairing load bearing bone defects [7, 8]. In vivo assessment of 13- 93 bioglass has shown good bonding between the scaffold and the surrounding tissues, in addition to new bone and tissue growth around the scaffold [9-11]. A 30% reduction in compression strength was reported upon immersion of the 13-93 bioglass scaffolds in simulated body fluid for four weeks primarily because of the conversion of the 13-93 bioglass into hydroxyapatite (HA) [12].

Though the degradation has reportedly slowed down after two weeks and in fact, a recent long term in vivo assessment by Yinan et al. [13] showed that about half ofthe scaffold still remains unconverted after 6 months of implantation in rat calvarial defects. Thus making the 13-93 bioglass a potential material to fabricate an implant for long term load bearing repair.Previous studies on hot pressing of bioglass-metal composites have shown that the flexural strength and toughness of composite improved upon dispersion of metallic particles in the glass matrix [13-16]. Unsuccessful attempts to produce such composites by pressure less sintering were previously reported in the literature which could have been failed due to differences in the thermal properties of the glass and the metal. Also, there have been no reports of manufacturing such composite parts for biomedical applications using extrusion based additive manufacturing processes.

Also, there have been no reports of manufacturing such composite parts for biomedical applications using extrusion based additive manufacturing processes. As to provide more structural integrity for the scaffolds, we investigate addition of ductile titanium (Ti) metallic fibers into the 13-93 bioglass matrix to form a composite with improved toughness. Ti is used due to the following reasons: 1) Ti is an FDA approved, and widely used material for synthetic implants as it offers excellent biocompatibility; 2) there is only a small difference in the thermal expansion coefficients of 13-93 bioglass and Ti fibers; and 3) the presence of silica in 13-93 bioglass matrix and titanium oxide on the surface of titanium fibers will essentially provide good adhesion between these materials, thereby reducing problems like de-bonding.

Therefore, the addition of metallic fibers is expected to improve the toughness of the implant and avoid a catastrophic failure during the bone repair process.Several additive manufacturing techniques have been used in the past to fabricate scaffolds. Robocasting was used by Cesarano et al. to fabricate lattice scaffolds for load-bearing bone repair [19]. Simpson et al. [20] and Kolan et al. [21, 22] studied the effect of pore sizes and pore geometry on mechanical and biological properties of scaffolds manufactured by selective laser sintering. Foam replication technique was used by Chen et al. [23] to fabricate highly porous (90%) scaffolds. However the above additive manufacturing techniques had low toughness compared to conventional methods like hot pressing.

[23] to fabricate highly porous (90%) scaffolds. However the above additive manufacturing techniques had low toughness compared to conventional methods like hot pressing. 3D printing of composites, especially, bioglass and metal composites, is relatively unexplored and acquiring further importance with the advent of several successful extrusion based techniques.the scaffolds fabricated usingThis study aims at fabricating Ti reinforced glass scaffolds using freeform extrusion. An aqueous paste mix of 13-93 bioglass, additives and Ti fibers were made for extrusion. Two sets of scaffolds, one with fiber (0.4 vol. %) and other one without fiber were fabricated and their heat treatment schedule was identified. Following this a comprehensive evaluation of its mechanical and biological properties was performed.2.

Following this a comprehensive evaluation of its mechanical and biological properties was performed.2. Materials and methods2.1 Fabrication of scaffoldsThe as-received water quenched 13-93 bioglass (Mo-Sci Corp, Rolla, MO) was crushed in a steel shatterbox (SPEX SamplePrep Crusher, Model 8500, Metuchen, NJ) and attrition- milled using de-ionized water for 3 h with ZrO2 as the grinding medium. The particle size distribution was measured using a laser diffraction-based particle size analyzer (Model LS 13 320, Beckman Coulter Inc., Fullerton, CA). To make an extrudable paste, additives were added to the glass and then wet ball milled together overnight. Following milling, the mixture was heated for about 50 min at 70°C with continuous stirring along with the simultaneous addition ofbinder (Methocel).

Following milling, the mixture was heated for about 50 min at 70°C with continuous stirring along with the simultaneous addition ofbinder (Methocel). For preparing a paste with fibers, the as-received long fibers were crushed using a porcelain mortar and pestle crusher with moderate force and then added to the stirring mix along with Methocel. This ensured uniform dispersion of fibers within the paste. Following this the paste was vacuum mixed for 4-5 min (WhipMix Vacuum Power Mixer Plus, WhipMix Corporation, Louisville, KY) to remove any air bubbles and then transferred into an air-tight container. The paste composition is provided in Table 1.Table 1. Titanium reinforced 13-93 bioglass paste compositionComponent13-93 Glass particles Darvan CMethocelTitanium fiber De-ionized waterConcentration (vol.

Titanium reinforced 13-93 bioglass paste compositionComponent13-93 Glass particles Darvan CMethocelTitanium fiber De-ionized waterConcentration (vol. %) 40 24.0 - 4.40.0 - 0.4 54ManufacturerMo-Sci Corp, Rolla, MO Vanderbilt Minerals LLC, Norwalk, CT Dow Chemical Company, Midland, MI Intramicron LLC, Auburn, AL -A three-syringe independent extruder developed at Missouri S&T is used for fabrication of scaffolds. It primarily consists of extrusion devices, a motion subsystem and a real-time control system. The printer has X, Y, and Z-axes motion capabilities controlled by three stepper motors (Empire Magnetics, Rohnert Park, CA). A ram extruder is used to provide the extrusion force. The paste in a syringe (60 cc plastic syringe) is extruded through a 1.19 mm nozzle (Nordson EFD, Westlake, OH) on to a hot plate set at 40°C. After the completion of a layer, the gantry moves up by the thickness of one layer. These steps are repeated until the entire part is formed.

After the completion of a layer, the gantry moves up by the thickness of one layer. These steps are repeated until the entire part is formed. A 40 x 36 x 6 mm3 scaffold with ~800 µm pore size and 0.7 mm layer height was fabricated.2.2 Post processing and characterization of scaffoldsThermogravimetric analysis (TGA) (NETZSCH thermal analyzer STA 409, Burlington, MA) was performed on the ‘green parts’ made, with and without Ti fiber, in order to identify the post processing schedule. The heating rate, temperature holds, and sintering temperature were determined from the TGA results. The pore sizes of the sintered scaffolds were measured using optical microscopy. The apparent porosity of the sintered scaffolds was measured using Archimedes method. X-ray diffraction (Philips X-Pert, Westborough, MA), run over 2θ range of 10° – 90° was used to identify the amorphous nature of sintered glass as well as the presence of Ti, using Cu Kα radiation (λ = 0.154056 nm).

Surface of Ti fibers was examined for presence of oxides using X-ray photoelectron spectrometer (Kratos Axis 165 Photoelectron Spectrometer, Manchester, UK).2.3 Mechanical testing of scaffoldsScaffolds (3 x 5 x 25 mm3) with different volume fractions of titanium fibers were tested in flexure using an Instron testing machine (model 5881, Norwood, MA). Prior to testing, surface grinding (FSG-618, Chevalier Machinery Inc., Santa Fe Springs, CA) was done to prepare parallel surfaces. A four point semi-articulated fixture (outer span of 20 mm and inner span of10 mm) at a cross-head speed of 0.2 mm/min using a 2 kN load cell was used in the flexure test and the load was applied along the z direction. The flexural stress was determined using the following equation (ASTM C1674-11):𝜎 =3𝑃𝑙 4𝑏𝑑2where P is the applied force, l is the length of the outer span, b is the width and d is the thickness of the sample.

The strength of the samples tested is expressed as mean ± standard deviation (SD).A Chevron notched beam test was used to assess the fracture toughness of the scaffolds (3 x 5 x 25 mm3). A notch was made at the mid-span of the scaffold using a dicing saw (Accu- Cut 5200, AREMCO Products Inc., Ossining, NY) with a 0.15 mm thick diamond blade. A four- point, semi-articulated fixture mentioned above was used for this testing too. The fracture toughness was calculated using the following equation (ASTM C1421-10):𝐾𝑐 = 𝐹𝑚𝑌∗𝑚𝑖𝑛(𝑆𝑜 − 𝑆𝑖)10−6/𝐵𝑊3/2where Kc is the fracture toughness, Fm is the maximum load, So and Si are outer and inner spans of the fixture used, B is the depth, and W the width of the specimen. Y* min is the minimum of the geometric function (calculated based on ASTM C1421-10 guidelines).

Y* min is the minimum of the geometric function (calculated based on ASTM C1421-10 guidelines). At least five samples were tested and the toughness is expressed as mean ± SD.2.4 SBF tests on scaffoldsFor in vitro assessment two sets of scaffolds were used, one set without fiber and the other set with fiber. The scaffolds (5 x 5 x 5 mm3) were cleaned thrice in distilled water, then in ethanol using an ultrasonic cleaner (Crest CP 500T, Trenton, NJ), and then dried overnight at 65℃. The scaffolds were then weighed and immersed in simulated body fluid (SBF) prepared according to the Kokubo method with a starting pH of 7.40. 100 ml of SBF solution was used per gram of the scaffold, and the scaffolds along with SBF solution were stored in plastic Nalgene bottles [24]. The samples were then kept in an incubator maintained at 37 ℃. Scaffolds were removed every week until four weeks and then dried overnight before weight and strength were assessed.

The samples were then kept in an incubator maintained at 37 ℃. Scaffolds were removed every week until four weeks and then dried overnight before weight and strength were assessed. Five samples per set were used in the compression testing and the strength is expressed as mean ± SD. Scanning electron microscopic (Hitachi S-4700 FESEM, Hitachi Co., Tokyo, Japan) images were taken to study the surface morphology of the scaffolds. The scaffolds removed from SBF were tested in compression using an Instron testing machine (model 5881, Norwood, MA) at a crosshead speed of 0.5 mm/min using a 10 kN load cell. At least five samples were used in the compressive strength measurements and results were reported as mean ± SD.3. Results and Discussion3.1 Fabrication of scaffoldsThe size distribution of attrition milled glass particles is shown in Figure 1.a. The particle size varied from ~0.2 to ~12 µm with an average of 2.3 µm.

The particle size varied from ~0.2 to ~12 µm with an average of 2.3 µm. The SEM image of the milled bioglass and the as-received titanium fibers are shown in Figure 1.b and 1.c, respectively. A typical irregular shape of the milled glass particles can be observed which will aid in achieving appropriate viscous flow characteristics during sintering. This could potentially reduce the micropores in the scaffolds [25]. The Ti fibers used in this study had an average diameter of 16 µm and aspect ratio of 250.Figure 1. (a) Particle size distribution of attrition milled bioglass (b) SEM image of the attrition milled 13-93 bioglass (c) SEM image of as-received titanium fibers.The fabrication process consists of three basic steps: preparing the paste, printing the scaffolds and sintering them. 13-93 bioglass+Ti composite paste was prepared with varying volume content of the Ti fibers from 0.1% to 0.5% in increments of 0.1%.

13-93 bioglass+Ti composite paste was prepared with varying volume content of the Ti fibers from 0.1% to 0.5% in increments of 0.1%. The volume content was kept at 0.4% as a further increase in the Ti fiber content consistently blocked the needle and caused difficulties during the fabrication. The prepared paste (Figure 2.a) is filled into a syringe (manually) and is extruded onto a hot plate maintained at 40℃ using layer-by-layer deposition through a 1.19 mm diameter nozzle (Nordson EFD, Ohio, USA). Scaffolds were fabricated using the extrusion fabrication machine shown in Figure 2.b. The extrusion force varied between 200- 320 N based on the amount of titanium fibers in the paste and the viscosity of the paste. An as- fabricated scaffold with a 0.4% volume fraction of fibers is depicted in Figure 2.c. The green scaffold was 40 mm x 36 mm x 6.0 mm in size and the pore sizes varied from 750 to 1000 µm.Figure 2.

The green scaffold was 40 mm x 36 mm x 6.0 mm in size and the pore sizes varied from 750 to 1000 µm.Figure 2. Steps in fabrication of scaffolds: (a) paste prepared with titanium fibers dispersed in bioglass; (b) Machine and syringe filled with paste used in fabrication of scaffolds; (c) As-fabricated scaffold; (d) Thermogravimetric analysis of paste used. The binder burnout schedule for the ‘green scaffold’ was developed from the TGA curve in Figure 2.d. This curve shows the changes in weight of the ‘green scaffold’ as a function of increasing temperature. Noticeable changes in the weight of the scaffold were observed atapproximately 100 °C and 300 ℃. The residual water evaporates at 80-120 ℃ and Methocel and Darvan C burn out at around 300-350 ℃. Holding temperatures were designed based on the curve, as to aid slow burn-out of the additives in the scaffold. These observations were used to prepare the binder burnout and sintering schedule in Table 2.Table 2.

These observations were used to prepare the binder burnout and sintering schedule in Table 2.Table 2. Binder burnout/sintering schedule used for scaffolds fabricatedStarting Temperature (℃) 20 350 700Ending Temperature (℃) 350 700 50Rate (℃/min)0.3 2 10Hold Period (h) 2 1 -A slower heating rate is adopted in the binder burnout process so as to allow sufficient time for the decomposition of the organic additives. Figure 3.a shows a scaffold that underwent the binder burnout process at a faster heating rate of 5 ºC/min (for 20 ºC to 350 º C segment). The part turned black. This is mostly due to the trapped carbon from the organic additives. A sintered scaffold with 0.4 vol. % fiber that underwent the binder burnout schedule in Table 2 is pictured in Figure 3.b. A slow heating rate provides sufficient time for the additives to decompose from the fabricated part.

% fiber that underwent the binder burnout schedule in Table 2 is pictured in Figure 3.b. A slow heating rate provides sufficient time for the additives to decompose from the fabricated part. The average shrinkage observed in the sintered scaffolds was measured to be ~25%, ~29% and ~17%, along the length, width, and thickness, respectively. The pore sizes varied from 600-850 µm. Optical images of the sintered scaffolds (Figure 3.c, d) reveal that the fibers are dispersed in the glass matrix. Figure 3.d shows the sintered part a higher magnification. The Ti fibers oriented along the direction of the extruded filament can also be clearly observed. The porosity of the scaffolds was measured to be ~58% based on the Archimedes method. Also, the XRD patterns of the sintered glass scaffold and the as-received glass particles indicated no crystallization and the glass retained its amorphous nature.abcdFigure 3.

Also, the XRD patterns of the sintered glass scaffold and the as-received glass particles indicated no crystallization and the glass retained its amorphous nature.abcdFigure 3. (a)Scaffold that turned black after undergoing binder burnout at a faster rate (5 ºC/min); (b) Sintered 13-93 bioglass scaffold with 0.4 vol. % titanium fibers; (c) & (d) Optical image of the scaffold in (b) with its zoomed version3.2 Assessment of mechanical properties of scaffolds3.2 Assessment of mechanical properties of scaffoldsThe flexural strength, flexural modulus and fracture toughness of scaffolds with varying vol. % of fibers is pictured in Figure 4. The flexural strength varied from ~10 MPa (0 vol. % of Ti fibers) to ~15 MPa (0.4 vol. % of Ti fibers). The flexural strength increased as the Ti fiber volume fraction increased.

The flexural strength varied from ~10 MPa (0 vol. % of Ti fibers) to ~15 MPa (0.4 vol. % of Ti fibers). The flexural strength increased as the Ti fiber volume fraction increased. Previous mechanical testing studies on 13-93 bioglass scaffolds prepared by robocasting have shown that the value of flexural strength of 13-93 bioglass scaffolds is in the range of 11±3 MPa which are consistent with the values obtained in this study [8]. The flexural modulus of the scaffolds increased to 15 GPa from 11 GPa due to addition of 0.4 vol. % Ti fibers into the scaffold. The trend in flexural modulus clearly shows that the flexural strength is increasing with addition of fibers. The flexural strength reported in this study is similar to that of human trabecular bone (10-20 MPa) [26, 27]. The scaffolds with 0.4 vol. % of Ti fibers were found to have the highest fracture toughness (~ 0.8 MPa•m1/2 ) and scaffolds without fibers had a toughness of ~ 0.5 MPa•m1/2.

The scaffolds with 0.4 vol. % of Ti fibers were found to have the highest fracture toughness (~ 0.8 MPa•m1/2 ) and scaffolds without fibers had a toughness of ~ 0.5 MPa•m1/2. This trend can be attributed to the presence of high amount of Ti fiber content in the scaffolds with 0.4 vol. % fibers. The toughness of the scaffolds with 0.4 vol. % fibers is comparable to that of human trabecular bone (0.1-0.8 MPa•m1/2 ) [26, 27]. Both flexural strength and fracture toughness improved with the addition of Ti fibers.Figure 4: (a) Flexural modulus and flexural modulus (b) fracture toughness of scaffold without (0 vol. %) and with Ti fibers (0.4 vol. %)Studies have shown that the mechanical properties of borosilicate glass matrix improved (five-fold improvement in fracture toughness) due to addition of Hastelloy X fibers (15 vol. %) in them [14]. XPS analysis performed on the fibers showed a presence of TiO2 on the surface of the Ti fibers.

%) in them [14]. XPS analysis performed on the fibers showed a presence of TiO2 on the surface of the Ti fibers. The presence of a layer of TiO2 could aid in bonding of the Ti fiber with the 13-93 bioglass matrix which predominantly consists of several oxides including SiO2. In addition, the rough and uneven surface of Ti fiber assists the bioglass matrix to better bond with the Ti fibers. An increased adhesion between the glass matrix and the Ti fibers will increase the strength of scaffolds [29]. Below a fine layer of TiO2 on surface of the fibers, pure elemental Ti is present at the core. The Ti fibers have an average tensile strength in the range from 246 to 370 MPa which would help reinforce the brittle glass matrix by transferring the bending stresses from the matrix to the fibers [14, 28].In this study, Ti fibers were added during the production of paste and not during the fabrication process.

This type of pre-impregnation creates a homogenous matrix, which in turn increases the strength considerably as compared to when the fibers are added manually. Figure 5 shows the SEM images of the cross section of sintered pellets which shows the Ti fiber bonded to the glass matrix. The thermal expansion coefficient of Ti fibers and 13-93 bioglass are 8.6 x 10-6 m/m °C and 10.6 x 10-6 m/m °C respectively. After sintering, the glass matrix will be in contraction and the Ti fibers will be in tension, preserving the bond between the matrix and fiber. No fiber pullout was observed after the sintering process. The fibers in scaffold are expected to prevent crack propagation and thereby toughens the 13-93 bioglass matrix. For the scaffolds without any fibers, the cracks could easily propagate through the brittle matrix leading to fracture.Figure 5. 13-93 bioglass-Ti fiber pellets showing good bonding between Ti fibers andglass matrix.3.3.

13-93 bioglass-Ti fiber pellets showing good bonding between Ti fibers andglass matrix.3.3. In-vitro evaluation of the scaffoldsMechanical degradation of scaffolds without and with Ti fibers, as a function of immersion time in SBF is shown in Figure 6.a and 6.b respectively. The compressive strength of scaffolds without fibers, before immersion, was measured 103 ± 33 MPa. This strength was reduced to 67 MPa (~ 30 % reduction) following the immersion of scaffolds in SBF for four weeks. The strength of scaffolds with fibers before immersion was 128 ± 30 MPa. This strength was reduced by ~39% after 4 weeks to 88 MPa. The compressive strength reduced considerably in the first two weeks (~24 %). It was also observed that the reduction in strength is low from two to four weeks in comparison to the first couple of weeks (~10%). Such a reduction in the compressive strength is in accordance with similar results reported by researchers [5, 7, 8].

Such a reduction in the compressive strength is in accordance with similar results reported by researchers [5, 7, 8]. The compressive strengths of scaffolds with Ti fibers obtained even after four weeks of immersion in SBF was similar to that of human cortical bone proving the potential of the 13-93 bioglass + Ti composite system for load-bearing bone repair [26, 27].Figure 6. Variation in the compressive strength of the scaffolds after immersion in SBF for 4weeks; (a) for scaffolds without fiber (b) scaffolds with fiberThe SEM images of scaffold surfaces, with and without fibers immersed in SBF for two weeks, are shown in Figure 7. Figure 7.a shows the SEM image of the surface of the dried scaffold without fibers, after it has been removed from the SBF and dried overnight. Surface of a scaffold with fibers immersed in SBF is pictured in Figure 7.b. No noticeable changes could be observed between the both sets of scaffolds.

Surface of a scaffold with fibers immersed in SBF is pictured in Figure 7.b. No noticeable changes could be observed between the both sets of scaffolds. The surface cracks are typically formed because of the dried silica rich layer beneath the crystalline HA-like layer formed on the surface. Needle- like structures similar to that of HA were observed at a higher magnification (Figure 7.b) [3, 8, 12]. The XRD patterns of scaffolds (both with and without fibers) immersed in SBF had peaks that matched to the reference synthetic HA [24].Figure 7. SEM images of surface of a typical scaffold (a) without fiber (b) with fiber after immersion in SBF for 2 weeks and drying; (c) needle like HA crystals formed on the surface (at higher magnification).The formation of HA based on the SEM images and XRD patterns confirms the bioactivity of the scaffolds. When the 13-93 bioglass scaffolds are kept in SBF, Na+, K+ and (SiO4)4- ions are released into the solution.

When the 13-93 bioglass scaffolds are kept in SBF, Na+, K+ and (SiO4)4- ions are released into the solution. HA is formed due to the reaction of CaO present in the bioglass with phosphate ions in the SBF solution. The conversion of bioglass scaffold into HA also results in the reduction of its mechanical properties. Though the rate of degradation of silicate based 13-93 bioglass is slower because of its chemical durability. This can be observed in the reduction in compressive strength of the scaffolds after four weeks of immersion in SBF; SeeFigure 5. Such a slow reduction in compressive strength of the scaffolds is a desirable property of bone implants. The scaffolds with fibers had an average compressive strength of ~88 MPa after immersion in SBF solution for four weeks. This finding leads to a possibility that these scaffolds could be used for load bearing applications in the body, as it is comparable to the that of human cortical bone (100 – 150 MPa) [26, 27].4.

This finding leads to a possibility that these scaffolds could be used for load bearing applications in the body, as it is comparable to the that of human cortical bone (100 – 150 MPa) [26, 27].4. ConclusionThis study investigated the feasibility of fabricating titanium fiber reinforced 13-93 bioglass scaffolds using freeform extrusion fabrication technique. Scaffolds reinforced with fibers had a fracture toughness of ~ 0.8 MPa•m1/2 and flexural strength of ~ 15 MPa. The fracture toughness of scaffolds with fibers increased by 70% compared to that without fibers and the flexural strength increased by 40%. The in vitro assessment of scaffolds revealed that the addition of biocompatible titanium fibers to the bioactive glass reinforced the scaffold mechanically without inhibiting its bioactive properties.

The improved mechanical properties with compressive strengths of ~88 MPa even after four week degradation in simulated body fluid shows the potential of the 13-93 bioglass+Ti composite implants for load bearing bone repair applications.5. AcknowledgmentThis research was funded by the Interdisciplinary Intercampus Research Program of University of Missouri System and the Center for Biomedical Science and Engineering at the Missouri University of Science and TechnologyReferences[1][2][3][4][5][6]Hench, L. L., Bioceramics: from concept to clinic. J. American Ceramic Society. 74(7), 1487-1510 (1991). Fu, Q., Rahaman, M. N., Bal, B. S., Huang, W., & Day, D. E., Preparation and bioactive characteristics of a porous 13–93 glass, and fabrication into the articulating surface of a proximal tibia. J. Biomed. Mat. Res. Part A. 82(1), 222-229 (2007). Liu, X., Rahaman, M. N., Fu, Q., & Tomsia, A.

eeform Extrusion of High Solids Loading Ceramic Slurries, Part I: Extrusion Process ModelingMichael S. Mason1, Tieshu Huang2, Robert G. Landers1, Ming C. Leu1, and Gregory E. Hilmas21870 Miner Circle Department of Mechanical and Aerospace Engineering1 Department of Materials Science and Engineering2 University of Missouri - Rolla, Rolla, Missouri 65409–0050 {mmason,hts,landersr,mleu,ghilmas}@umr.eduReviewed, accepted September 14, 2006ABSTRACTA novel solid freeform fabrication method has been developed for the manufacture ofceramic-based components in an environmentally friendly fashion. The method is based on the extrusion of ceramic slurries using water as the binding media. Aluminum oxide (Al2O3) is currently being used as the part material and solids loading as high as 60 vol. % has been achieved. This paper describes a manufacturing machine that has been developed for the extrusion of high solids loading ceramic slurries.

% has been achieved. This paper describes a manufacturing machine that has been developed for the extrusion of high solids loading ceramic slurries. A critical component of the machine is the deposition system, which consists of a syringe, a plunger, a ram actuated by a motor that forces the plunger down to extrude material, and a load cell to measure the extrusion force. An empirical, dynamic model of the ceramic extrusion process, where the input is the commanded ram velocity and the output is the extrusion force, is developed. Several experiments are conducted and empirical modeling techniques are utilized to construct the dynamic model. The results demonstrate that the ceramic extrusion process has a very slow dynamic response, as compared to other non-compressible fluids such as water.

The results demonstrate that the ceramic extrusion process has a very slow dynamic response, as compared to other non-compressible fluids such as water. A substantial amount of variation exists in the ceramic extrusion process, most notably in the transient dynamics, and a constant ram velocity may either produce a relatively constant steady-state extrusion force or it may cause the extrusion force to steadily increase until the ram motor skips. The ceramic extrusion process is also subjected to significant disturbances such as air bubble release, which causes a dramatic decrease in the extrusion force, and nozzle clogging, which causes the extrusion force to slowly increase until the clog is released or the ram motor skips.1. INTRODUCTIONThere have been only a handful of novel SFF systems described in the literature over thepast several years [1-7]. Most of these systems have utilized 3D gantry positioning systems.

INTRODUCTIONThere have been only a handful of novel SFF systems described in the literature over thepast several years [1-7]. Most of these systems have utilized 3D gantry positioning systems. Malone [2] explained that 3D gantries have easy maneuverability and high payload capacities. The emphasis in this work was on high motor accelerations so that a constant deposition feed rate could be utilized. Weiss and Prinz [4] also developed this type of system for customized solid freeform fabrication processes. Some improvements have also been made to the Selective Laser Sintering (SLS) process by reducing the amount of material being heated, thereby improving part accuracy [5-6]. Other areas of improvement have been in Stereolithography by reducing the part shrinkage that is inherent in the process [7]. Freeze-form Extrusion Fabrication (FEF) is a new SFF process that deposits a ceramic slurry utilizing extrusion in a layer-by-layer316fashion to create 3D geometries [8].

Freeze-form Extrusion Fabrication (FEF) is a new SFF process that deposits a ceramic slurry utilizing extrusion in a layer-by-layer316fashion to create 3D geometries [8]. The FEF system uses a different approach in mechanical implementation as compared to most other SFF processes. Similar to Rapid Freeze Prototyping [9-12] an XY table with a separate Z elevator axis is used. The 3D gantry and the XY table with Z elevator operate on the same principle. The 3D gantry has all three axes connected into a single unit, which forces the axes to be level with respect to each other, within a specified tolerance, before installation. The XY table and Z elevator are individually known to be level, but the axes must be adjusted to be level relative to each other during installation.A large amount of ceramic flow modeling research has been conducted.

Benbow and Bridgewater developed the standard model for this process [13], which is a static relationship between ram velocity and extrusion force. Other modifications of this model have been developed [14], however this model is still static. A dynamic model of the extrusion process will be developed in this paper to better understand the process.In the following sections the FEF system will be described in detail, highlighting the major process parameters. The dynamics of the ceramic high solids loading slurry extrusion process will be examined, including system disturbances. In section two each piece of hardware comprising the FEF system is described, as well as its function and interactions with other hardware components. In section three the extrusion process is described. First, the internal system disturbances are discussed along with the effects they have on the extrusion process.

In section three the extrusion process is described. First, the internal system disturbances are discussed along with the effects they have on the extrusion process. Then experimental tests are conducted to characterize the dynamic behavior of the extrusion process.2. EXPERIMENTAL SETUPThe FEF experimental system (see Figure 1) is composed of three major components:motion system, real-time data acquisition and control system, and extrusion device. Each of the components can be decomposed into several subcomponents, which are described below.Three Empire Magnetic extended temperature range stepper motors are used to driveParker Hannifin Daedal 404 XR series linear axes. The three orthogonal linear axes each have 254 mm of travel, forming a three-dimensional motion system. The stepper motors have a stepping angle of 1.8 degrees, but have the ability to utilize “micro steps.” Micro steps give a motor the ability to take multiple smaller steps to reach a single stepping angle.

The stepper motors can vary the micro steps from 1-250 microsteps/full step giving a maximum possible resolution of 0.0072 degrees/step. The resolution of each axis is 2.5 (cid:80)m per step, or 0.1 (cid:80)m per micro-step, when using 250 microsteps/full step. The motors have a maximum rotational speed of 50 rev/s, which provides a maximum velocity of 250 mm/s for the individual axes. The maximum motor angular acceleration is 50 rev/s2, which provides a maximum linear axis acceleration of 250 mm/s2.An extrusion device capable of exerting high pressures is necessary for deposition of highsolids loaded ceramics. A stepper motor with a resolution of 16,000 steps/inch and maximum velocity of 15 inches/s is directly coupled to the Z-axis. The extrusion axis, having eight inches of stroke is used to move a plunging device, which applies pressure to a syringe to extrude ceramic material. The deposition rate can be changed online by adjusting the commanded stepper motor speed.

The deposition rate can be changed online by adjusting the commanded stepper motor speed. A 60 cm3 syringe is used to hold and eject the deposition material. Nozzles of diameters ranging from 190 – 580 (cid:80)m are attached to the end of the syringe to allow for different thicknesses of material to be deposited.317A LC-305 load cell from Omega Engineering is utilized for extrusion force feedback. The load cell measures the force the ram extruder applies to the ceramic slurry. The load cell outputs a differential voltage (0 – 20 mV) that corresponds to the extruded force (0 – 4450 N). This signal is sent to an amplifier that increases the differential voltage by a factor of 100, thus, the output range is 0 – 2 V. This increases the input range, thereby increasing the overall resolution of the feedback device.

This increases the input range, thereby increasing the overall resolution of the feedback device. The load cell has a 0.1% linearity, which gives an output within +/- 4.45 N of the actual force.Extrusion DeviceZ AxisLoad CellX AxisY AxisStepper MotorsXY TableFigure 1: The FEF experimental system.3. FLOWRATE MODELINGIn modeling and controlling the ceramic slurry extrusion process, one must consider the system disturbances and material flow dynamics. The first subsection below describes in detail the two major disturbances that are present within the extrusion process: agglomerate breakdown and air bubble release. The next subsection explains the different types of material flowrate tests conducted in order to characterize the dynamic extrusion process.3.1 System DisturbancesThere are two major disturbances that occur during the extrusion process: agglomeratebreakdown and air bubble release.

Both of these disturbances cause changes to the ceramic slurry flowrate. The following sections explain why the disturbances are present and how they affect the extrusion process.3.1.1 Agglomerate Breakdown Ceramic slurries are fluids with significant solids loading. In the case of the FEF process, the solids loading is higher, generally greater than 50 vol. %. The solid particles that are present within the slurry are referred to as agglomerates. When batching the ceramic slurry the ceramic powder is mixed with an aqueous media. The current ceramic of interest is alumina (Al2O3), in combination with water as the binding media, Polyethelene-Glycol for improved green part strength, and Aquazol for material lubrication. The materials are mixed318for a period of time to form an isotropic slurry. The ceramic particle size is reduced as much as possible, generally ranging from 0.8 to 2.5 (cid:80)m.

The materials are mixed318for a period of time to form an isotropic slurry. The ceramic particle size is reduced as much as possible, generally ranging from 0.8 to 2.5 (cid:80)m. With particles of varying size it becomes increasingly difficult for the particles to be evenly interspersed within the aqueous media. The ceramic particles join together into larger particles called agglomerates. The agglomerates are interspersed within the aqueous media, thereby forming the ceramic slurry.As force is applied to the ceramic slurry, the material reservoir compresses due to the presence of air bubbles (described later), and material flowrate increases with the rising force. After compression is complete the material flowrate reaches a stable state. As the material enters the nozzle, the space available for agglomerates to be interspersed is significantly reduced. This causes an increase in resistance force for compression of the slurry since the agglomerates are pressed together.

This causes an increase in resistance force for compression of the slurry since the agglomerates are pressed together. After a critical force is reached the agglomerates “slide” across edges of other agglomerates. During the sliding process the agglomerates break into smaller agglomerates allowing for material flow to continue. Figure 2 shows an example of agglomerate breakdown.The force spikes caused by agglomerate breakdown cannot be reliably predicted. The agglomerate size is random and is different for each slurry preparation. The breakdown of the agglomerates also occurs randomly, but can be reduced by minimizing the ceramic particle size.Figure 2: Illustration of agglomerate breakdown.3.1.2 Entrapped Air Bubble Release During the preparation phase of the ceramic slurry, the slurry is mixed in a vacuum to remove air present within the slurry for a more consistent fluid. After the slurry is mixed, it must be loaded into the fluid reservoir.

After the slurry is mixed, it must be loaded into the fluid reservoir. Due to the high solids loading the fluid is extremely viscous and requires manual loading into the syringe. During the319loading process air becomes entrapped, forming bubbles within the ceramic slurry. As the ram applies force to the ceramic slurry, the air bubbles compress due to buoyancy effects from the surrounding slurry. As material is extruded the air bubbles are forced into the die region (see Figure 3), which is much smaller than the rest of the syringe. The air bubbles join together as they enter the die region, forming larger bubbles (Figure 4). As the air bubbles reach the nozzle exit they pop, creating a large pressure release. This causes the extrusion force to dramatically drop.

As the air bubbles reach the nozzle exit they pop, creating a large pressure release. This causes the extrusion force to dramatically drop. The region previously occupied by the air bubble is now empty allowing for the ceramic slurry to expand.Figure 3: Illustration of slurry flow within material reservoir.3.2 Flowrate TestsFor modeling the ram extrusion process two different types of tests were conducted. Itwas first necessary to determine the ram velocity operational range. This was determined by running the ram extrusion velocity at various speeds and visually inspecting the extrudate for curling and clogging. Curling is the phenomenon where the extrudate curls on the nozzle tip, as shown in Figure 5. Clogging is when extrudate dries within the nozzle and reduces the material flowrate until no material can be extruded through the nozzle. Neither of these situations are desirable.

Clogging is when extrudate dries within the nozzle and reduces the material flowrate until no material can be extruded through the nozzle. Neither of these situations are desirable. The extrusion velocity range for the ceramic slurries in this paper was determined to be 0.2 to 0.65 (cid:80)m/s. Velocities below the lower range lead to clogging, while velocities above the higher range cause curling, which will eventually lead to clogging. It should be noted that velocities outside the operational range can be used, but only for short periods of time.320Applied ForceAir bubbles - OAir bubbles joiningFigure 4: Illustration of entrapped air bubbles within material reservoir.Figure 5: Curling of extrudate on nozzle tip.3.2.1 Long-Step Tests A series of step tests with constant ram velocities were conducted. The commanded ram velocity was kept constant for one hour and extrusion force and mass data were collected at a rate of 1 Hz.

The commanded ram velocity was kept constant for one hour and extrusion force and mass data were collected at a rate of 1 Hz. The mass data was collected through a mass balance connected to the PC through a USB port. The mass balance was set to a baud rate of 9200 and continuous acquisition. A total of seven tests were conducted at different velocities within the range of interest using a 190 (cid:80)m diameter nozzle. During most of the tests the extrusion force rose exponentially during the transient phase and a relatively constant force was reached in the steady state. However, in some tests the force continued to rise at a slow rate during the entire test, never reaching a relatively constant force.Empirical modeling was performed to determine a model for each test conducted. Afirst-order model of the formK s (cid:87) (cid:14) where K is the gain and (cid:87) is the time constant. This model was fit to each set of experimental data.

Afirst-order model of the formK s (cid:87) (cid:14) where K is the gain and (cid:87) is the time constant. This model was fit to each set of experimental data. To find the steady-state extrusion force an average was taken over the region following thesF )( )( sv(cid:32)1321(1)initial force rise. The time constant varied for each of the extrusion velocities, but each was found to be at a minimum 200 s. Table 1 shows the ram velocity, time constant, gain, and steady state extrusion force for each test. Figures 6-8 show some of the test and model results. Figures 7 and 8 show that the process may not be repeatable. For the same velocity input, test 2 reached a steady-state extrusion force in approximately 15 min; however, during test 3 the extrusion force increased during the entire testing period (i.e., 60 min).

The larger response time may be due to the viscosity characteristics of the ceramic slurry. If the slurry is a non-Newtonian fluid, then the viscosity increases with increasing fluid shear rate, which is directly affected by the ram velocity.

If the slurry is a non-Newtonian fluid, then the viscosity increases with increasing fluid shear rate, which is directly affected by the ram velocity. Future work will need to be conducted to investigate this hypothesis.500400300)N( e c r o F2001000Model Actual100 0102030 Time (min)405060Figure 6: Constant velocity test (#8) with ram extrusion v = 0.6 (cid:80)m/s and Fss = 320 N.322200150)N( e c r o F100Model Actual500Nozzle Nozzle Clogs Clogs501020 Time (min)3040Figure 7: Constant velocity test (#2) with ram extrusion v = 0.4 (cid:80)m/s and Fss = 180 N.250200)N150( e c r o F100500 0102030 Time (min)405060Figure 8: Constant velocity test (#3) with ram extrusion v = 0.4 (cid:80)m/s.Least squares was utilized to relate the three parameters, steady state extrusion force, ram velocity, and mass flowrate, to each other.

Three relationships were considered: mass flowrate versus extrusion force, mass flowrate versus ram velocity, and extrusion force versus ram velocity. Equations (2)-(3) show the relationships between the three parameters (mass flowrate, extrusion force, and ram velocity) and Figures 9 and 10 show the test data with the linear relationships. Even though the y intercept values were determined there is no physical meaning to them because equations (2)–(3) are only valid for ram velocities in the range of 0.3-0.6 (cid:80)m/s. The (cid:39)‘s in (2)-(3) refer to the changes between the parameters in relation to each other over the given range of ram velocity values.m (cid:39) F (cid:39) ss m (cid:39) v (cid:39)ssss70.0(cid:32)50(cid:32)ss323(2)(3)6055Measured Model) s / g m50( e t a r45w o l f40s s a M353025 100200 Force (N) Figure 9: Steady-state mass flowrate versus steady-state extrusion force.

Error bars correspond to one standard deviation.1502503006055Measured Model) s / g m50( e t a r45w o l f40s s a M3530250.5 0.4 Velocity ((cid:80)m/s) Figure 10: Steady-state mass flowrate versus steady-state ram velocity. Error bars correspond to one standard deviation.0.20.30.60.7One standard deviation of the curve fit is shown by the error bars in Figures 9-10. Thestandard deviations are 3.4 mg/s and 2.1 mg/s for equations (2)-(3), respectively. As can be seen from the figures, there is a definite correlation relating the three parameters. The correlation coefficients for mass flowrate to extrusion force, mass flowrate to ram velocity, and extrusion force to velocity are 0.9356 and 0.9764 respectively.3.2.2 Short-Step Test The next test conducted was a short step test in which the ram extruder ran at velocities covering the range of interest. The velocities incremented and decremented by 0.05 (cid:80)m/s every 40 seconds.

The velocities incremented and decremented by 0.05 (cid:80)m/s every 40 seconds. The initial velocity was run for a period of 8 min, which was determined to be the average time for a given ram velocity to reach a steady-state extrusion force, in order for the extrusion force to be constant before the test was started. The time history of the commanded velocity and corresponding extrusion force are shown in Figure 11. Force data was collected at a rate of 20 Hz.324) s / m (cid:80) (0.6y t i c o e Vl0.40.20200400600250)N200( e c r o F150100 0300 Time (s) Figure 11: Ram velocity and extrusion force time history for short step test.100200400500600As discussed before, the extrusion force increases with an increase in ram velocity. After examining the extrusion force data it can be seen there is a delayed reaction between the change in ram velocity and the change in the extrusion force.

After examining the extrusion force data it can be seen there is a delayed reaction between the change in ram velocity and the change in the extrusion force. The ram velocity decreases by 0.05 (cid:80)m/s for the first 5 time periods, but the extrusion force does not start to decrease until 200 sec, at which point the ram velocity has dropped to 0.2 (cid:80)m/s. At 240 sec the ram velocity is increased to 0.25 (cid:80)m/s, but the extrusion force continues to decrease until 280 sec. At this time the ram velocity is 0.3 (cid:80)m/s. The ram velocity continues to increase up to 0.6 (cid:80)m/s at 520 sec while the extrusion force keeps increasing until 640 sec, two time periods after the ram velocity has been decreased. During the remaining time the extrusion force is constant. Another interesting trend that can be seen in Figure 11 is that the extrusion force increases at a faster rate with increasing ram velocity than it decreases with decreasing ram velocity.

Another interesting trend that can be seen in Figure 11 is that the extrusion force increases at a faster rate with increasing ram velocity than it decreases with decreasing ram velocity. This is believed to be due to material compression effects. The ceramic slurry is a solid/liquid mixture. Both of these are considered incompressible material states, but with the presence of agglomerates and air bubbles within the material reservoir a small amount of compression is possible. As previously mentioned the agglomerates can breakdown which allows for compression as well as the air bubbles present. As material within the reservoir reduces, the amount of compression possible decreases due to the extrusion of air bubbles and the breakdown of agglomerates to a relatively small size.Another type of short-step test consisted of applying a constant ram velocity for a fixedperiod of time and then reducing the ram velocity for the same period of time.

This ram velocity cycle is repeated for the entire test.Two different types of impact tests were conducted. Extrusion force data was collected at a rate of 10 Hz. The first test conducted had lower ram velocities of 0.5 (cid:80)m/s and higher ram velocities that started at 5 (cid:80)m/s and increased by 5 (cid:80)m/s every other time period. Figure 12 shows the results. As can be seen the higher the ram velocity the faster the extrusion force increased. Also, there is a delayed reaction between the change in ram velocity and change in extrusion force, which ranges from 0-9 sec. It can be seen that the extrusion force increased so fast the ram motor eventually skipped.3252000) s / m (cid:80) ( 0 0 1 * v15001000Velocity ForceMotor skips, )N( F5000 020406080100Time (s) Figure 12: Impact test with increasing ram velocities.The second test conducted repeatedly changed the ram velocity between 0.5 and 5 (cid:80)m/s every 40 sec.

The application periods were too long, thereby causing the motor to skip as shown in Figure 13. The time period was then reduced to 10 sec for the last two test runs. Figure 14 shows more clearly the effect the value of the extrusion force has on the rate at which the extrusion force increases and decreases. The 0-9 sec delay is present for every change in ram velocity input. With a total of 73 individual ram velocity changes it was visually examined that the delay for increasing ram velocity was higher than for decreasing ram velocity. The average delay for increasing ram velocity was 4.3 sec with a standard deviation of 1.31 sec as calculated from 37 ram velocity increases.

The average delay for increasing ram velocity was 4.3 sec with a standard deviation of 1.31 sec as calculated from 37 ram velocity increases. The average delay for decreasing ram velocity was 2.5 sec with a standard deviation of 1.5 sec calculated from 36 ram velocity decreases.2000Velocity Force) s / m (cid:80) ( 0 0 1 * v, )N( F15001000500Motor skips0 0400 Time (s) Figure 13: Repeated ram velocity inputs for 40 s periods.2006008003262000) s / m (cid:80) ( 0 0 1 * v15001000Velocity Force, )N( F5000 050100 Time (s)150200Figure 14: Repeated ram velocity inputs for 10 s periods.4. SUMMARY AND CONCLUSIONS A new environmentally friendly solid freeform fabrication technique, Freeze-form Extrusion Fabrication (FEF), has been developed utilizing aqueous ceramic slurries as the build material. The FEF system was described in detail highlighting the major hardware components.

The FEF system was described in detail highlighting the major hardware components. System disturbances were described along with their effects on material flowrate behavior. Multiple tests were conducted in order to characterize the dynamics of the ceramic slurry utilizing a ram extruder. A first-order process model was developed and compared to collected data. From the large time constant values obtained from the extrusion tests conducted it wasshown that the slurry flowrate is much slower when compared to other incompressible fluid flows such as water. There is a large variability in the value of the time constant, indicating that the process is not very repeatable. The ram velocity, extrusion force, and mass flowrate have also been shown to be highly correlated from the long-step tests conducted with correlation coefficients of 0.93 or higher.

The ram velocity, extrusion force, and mass flowrate have also been shown to be highly correlated from the long-step tests conducted with correlation coefficients of 0.93 or higher. From the short-step tests it was shown that there are likely to be different dynamics present between increasing and decreasing extrusion force because there is a faster increase in extrusion force from higher ram velocities than a decrease in extrusion force due to a reduction in ram velocity.ACKNOWLEDGEMENTSThis work was supported by the Air Force Research Laboratory under Contract FA8650-04-C-5704. The authors would also like to thank Mike Hayes of Boeing for his technical support and Justin Kayser for his time and effort in this research project.REFERENCES 1. K. Taminger, R. Hafley, D. Fahringer, and E. Martin, “Effect of Surface Treatments on Electron Beam Freeform Fabricated Aluminum Structures,” Langley Technical Reports (2004).2. E. Malone, R. Rasa, D. Cohen, T. Isaacson, H.

eeform Extrusion of High Solids Loading Ceramic Slurries, Part II: Extrusion Process ControlMichael S. Mason1, Tieshu Huang2, Robert G. Landers1, Ming C. Leu1, and Gregory E. Hilmas21870 Miner Circle Department of Mechanical and Aerospace Engineering1 Department of Materials Science and Engineering2 University of Missouri at Rolla, Rolla, Missouri 65409–0050 {mmason,hts,landersr,mleu,ghilmas}@umr.eduReviewed, accepted September 14, 2006ABSTRACT Part I of this paper provided a detailed description of a novel fabrication machine for high solids loading ceramic slurry extrusion and presented an empirical model of the ceramic extrusion process, with ram velocity as the input and extrusion force as the output. A constant force is desirable in freeform extrusion processes as it correlates with a constant material deposition rate and, thus, good part quality. The experimental results in Part I demonstrated that a constant ram velocity will produce a transient extrusion force.

The experimental results in Part I demonstrated that a constant ram velocity will produce a transient extrusion force. In some instances the extrusion force increased until ram motor skipping occurred. Further, process disturbances, such as air bubble release and nozzle clogging that cause sudden changes in extrusion force, were often present. In this paper a feedback controller for the ceramic extrusion process is designed and experimentally implemented. The controller intelligently adjusts the ram motor velocity to maintain a constant extrusion force. Since there is tremendous variability in the extrusion process characteristics, an on-off controller is utilized in this paper. Comparisons are made between parts fabricated with and without the feedback control. It is demonstrated that the use of the feedback control reduces the effect of process disturbances (i.e., air bubble release and nozzle clogging) and dramatically improves part quality.1.

It is demonstrated that the use of the feedback control reduces the effect of process disturbances (i.e., air bubble release and nozzle clogging) and dramatically improves part quality.1. INTRODUCTION As described in Part I of this paper the material flowrate of high solids loading ceramicslurry extrusion is slow when compared to other incompressible fluid flows such as water [1]. Therefore a long period of time is required to reach the desired extrusion force. Applying a high ram velocity increases the extrusion force rapidly; however, due to compression effects, the extrusion force increases at an exponential rate eventually causing the motor to “skip.” This is an undesirable effect due to inconsistent deposition and possible ram motor damage. In this study a ram motor that can apply forces up to 2.2 kN before skipping occurs was utilized.

In this study a ram motor that can apply forces up to 2.2 kN before skipping occurs was utilized. An encoder, which is integrated into the motor, allows for position and velocity feedback data.From the literature reviewed by the authors, a feedback control system has never beenimplemented in a ceramic extrusion process during the deposition stage. Feedback control was implemented prior to deposition during the mixing stage, by adding organic binder, acidic, or basic chemicals to change the ceramic slurry properties such as density and viscosity, thereby modifying deposition results [2]. Russell [3] did a significant amount of testing examining effects of process parameters and applying statistical process control (SPC) with plans to329implement it towards feedback control. Pressure predication has also been an area of interest towards the future implementation of feedback control.

Pressure predication has also been an area of interest towards the future implementation of feedback control. Application of neural networks has been applied for a range of materials in order to produce a good model predictor [4].The authors have taken a different approach towards feedback control of the ceramicextrusion process. In order to obtain good material deposition and avoid a large period of time necessary to reach the steady-state extrusion force, an on/off controller has been implemented into the Freeze-form Extrusion Fabrication (FEF) system. Deposition tests are conducted and force data is analyzed for effects of constant ram velocity versus variable ram velocity. Parts that have been fabricated with and without the extrusion force controller are examined, comparing material deposition and surface roughness.2. EXTRUSION FORCE CONTROLLER DESIGNThe goal for material deposition in high solids loading ceramic extrusion processes is forsmooth, constant material flow.

EXTRUSION FORCE CONTROLLER DESIGNThe goal for material deposition in high solids loading ceramic extrusion processes is forsmooth, constant material flow. Material flowrate was determined to be directly related to the extrusion force (see Part I of this paper). For a constant ram velocity, the flowrate is transient. To achieve more consistent material extrusion a flowrate controller was designed and implemented. The feedback controller uses the extrusion force reading from a load cell to automatically adjust the ram velocity and maintain a constant extrusion force.The flowrate response of a high solids loading ceramic extrusion process is slow. It isundesirable to wait for the system to reach a steady-state extrusion force by applying a low ram velocity. Since extrusion force is directly related to ram velocity, a larger ram velocity will increase the extrusion force.

Since extrusion force is directly related to ram velocity, a larger ram velocity will increase the extrusion force. However, the ram extrusion motor can only apply forces up to 2.2 kN before it “skips” in order to reduce the torque being applied by the motor. After skipping, there is a drastic decrease in extrusion force, and a period of time is required to recompress the material and achieve the desired extrusion force. Due to this hardware limitation, a force limit must be set where the ram velocity is reduced before the motor reaches the skipping level.An on/off controller was implemented in order to maintain a constant extrusion force.The controller adds an extra initial level for feedback checks as compared with just two levels in a regular on/off controller. Three force levels and three ram velocities are used in the control algorithm. Figure 1 illustrates the on/off control algorithm.

Three force levels and three ram velocities are used in the control algorithm. Figure 1 illustrates the on/off control algorithm. The three levels of the force controller are ramp force Fr, lower-bound force Fl, and upper-bound force Fu. The ram velocities are ramp velocity vr, lower-bound velocity vl, and upper-bound velocity vh. The ramp force (Fr) is fixed during operation at a level approximately 90 N below the desired force (Fd), but can be changed to different values. Deposition during this time period is inconsistent and slow when compared to deposition when the extrusion force is at the desired value. Section I of Figure 1 shows the ramp force (Fr) and the ramp velocity (vr). The extrusion process begins with the ram velocity set to vr. After the ramp force is reached the extrusion force is still not at the desired level; however, it is still undesirable to operate the ram at a constant velocity to achieve the desired extrusion force.

An extrusion force range (i.e., upper and lower force bounds) is chosen that corresponds to an acceptable flowrate variation. This level is set to +/- (cid:71) N of the desired force. This force range can be changed, which is necessary due to variations in slurry viscosity from batch to batch. Due to variations in slurry preparation by different individuals there is a variation in slurry viscosity from “thinner” to “thicker.” The thinner slurries should have a lower (cid:71) value due to a faster response to the ram velocity. The thicker slurry tends to flow at a slower rate thereby reacting to the ram velocity at a slower rate allowing for a larger (cid:71) value. The lower-bound ram velocity (vl), which is slower than the ramp velocity, is applied330when the ramp force (Fr) is reached. This lower velocity allows for the ram to reach the desired force without causing the ram motor to skip.

This lower velocity allows for the ram to reach the desired force without causing the ram motor to skip. The implementation of these force limits and varying ram velocities reduce the time required to reach steady-state to 15–30 seconds, depending on the desired force, as compared to 13–60 minutes necessary when using a constant ram velocity.The last two ranges of operation are the acceptable material deposition range and theupper range in which the material flowrate is too fast for good deposition. The acceptable deposition range is when the extrusion force is between the lower and upper limits. If the extrusion force goes above the upper force (Fu), the ram velocity changes to vh in order to reduce the extrusion force, by allowing the material to decompress while continuing to flow through the nozzle. Depending upon several factors such as slurry viscosity and desired force, vh does not have to be 0.

Depending upon several factors such as slurry viscosity and desired force, vh does not have to be 0. As was shown in Part I, depending upon the desired extrusion force a value of 0 for vh may cause the extrusion force to decrease too quickly. In situations such as this a higher vh value is chosen so that the extrusion force decreases at a slower rate. Sections III-V in Figure 1 show how the controller chatters between the upper and lower bounds of the extrusion force. There is a delay in the flowrate due to the slow system response to ram velocity inputs.Figure 1: Flowrate controller force levels with corresponding ram velocities.The controller operates at a user defined rate, but is generally chosen to be 5 Hz. This is faster than the system is able to react, as previously discussed. This rate was chosen to speed up the reaction to the internal system disturbances (i.e., air bubble release and agglomerate breakdown).

This rate was chosen to speed up the reaction to the internal system disturbances (i.e., air bubble release and agglomerate breakdown). Depending on the nozzle diameter, the air bubble release can cause large force reductions between 25 N for the 190 (cid:80)m diameter nozzle and up to 450 N for the 580 (cid:80)m331diameter nozzle. The agglomerate breakdown is on a smaller scale and is independent of the nozzle diameter.Two other inconsistencies occur during material deposition, underfilling and overfilling. Underfilling is internal voids or gaps between points of material deposition. Overfilling is when excess material is deposited. Both of these are undesirable as they produce poor part quality either by creating unacceptable part porosity or poor surface finish. An additional problem with underfilling and overfilling is that they both lead to further problems as deposition continues.

An additional problem with underfilling and overfilling is that they both lead to further problems as deposition continues. With underfilling, the gaps that are created lead to gaps in the following layers that are above the gaps. This leads to overfilling as the material that is not deposited builds up on the nozzle and attaches to the next point where a gap in deposition is not present on the previous layer. Overfilling that occurs at any point during deposition will lead to material buildup on the nozzle during deposition of future layers. With the excess material present on the nozzle, material previously deposited attaches during motion over previous layers. This causes gaps in previously deposited layers. Practical examples of both underfilling and overfilling will be shown later.3. RESULTS To verify that the described extrusion force controller improves material deposition,several single-line material deposits were made.

RESULTS To verify that the described extrusion force controller improves material deposition,several single-line material deposits were made. The lines were 50 mm in length, using a table speed (vt) of 25 mm/s and a 580 (cid:80)m diameter nozzle. The first set of lines were deposited using a constant ram velocity (Figure 2). The second set of lines were deposited using the extrusion force controller (Figure 3).Beginning DepositionEnding DepositionFigure 2: Deposition lines using constant ram velocity (v = 0.5 (cid:80)m/s). Table speed is 25 mm/s and nozzle diameter is 580 (cid:80)m.Beginning DepositionEnding DepositionFigure 3: Deposition lines using the extrusion force controller. Table speed is 25 mm/s and nozzle diameter is 580 (cid:80)m.332By examining Figure 2 it can be seen that the material does not deposit in a constant fashion. The width of the lines varies significantly.

The width of the lines varies significantly. Figure 4 shows the ram velocity and extrusion force time history corresponding to the lines deposited in Figure 2. Although the ram velocity is constant the extrusion force continually changes during the entire extrusion process, causing inconsistent line width.250200Measured Desired)N( e c r o F15010050050100150 Time (s)200250300Figure 4: Extrusion force time history for lines deposited in Figure 2 (v = 0.5 (cid:80)m/s). Vertical lines indicate the beginning and ending of line deposition shown in Figure 2.By repeating the same deposition test but using the flowrate controller it can be seen from Figure 3 that the material deposition is much more consistent. In order to keep the test comparison as objective as possible the same environmental conditions (temperature, humidity, etc.) were maintained in both experiments as well as nearly the same amount of ceramic slurry left in the material reservoir.

A reference force of 613 N and a (cid:71) value of 0 were given for the on/off controller. The reference force of 613 N was chosen to show the ability of the controller to reach a desired extrusion force rapidly as compared to the application of a constant ram velocity. The (cid:71) value of 0 was chosen because it causes the most rapid reaction to a change in the extrusion force, thereby keeping the most constant extrusion force. Figures 5 and 6 show the ram velocity and extrusion force time history corresponding to the lines deposited in Figure 3. It should be noted that the extrusion force controller commands a ram velocity based on the extrusion force measured from the previous time period.

It should be noted that the extrusion force controller commands a ram velocity based on the extrusion force measured from the previous time period. Figure 6 shows this more clearly by zooming in on the portion between 114 and 120 seconds.3336) s / m (cid:80) (4y t i c o e Vl20 050100150200800Actual force Desired force)600N( e c r o F400200Fr0 050100 Time (s)150200Figure 5: Ram velocity and extrusion force time history for lines deposited in Figure 3. Vertical lines indicate the beginning and ending of line deposition shown in Figure 3.2) s / m (cid:80) (1.5y t i c o e Vl10.5618115116117118119 Actual force Desired force616)N( e c r o F614612610115116117 Time (s)118119Figure 6: Zoomed in portion of Figure 5.In another experiment, two bars were fabricated, one with a constant ram velocity and theother with the extrusion force control.

Figures 7 and 8 show pictures of the fabricated bars. As can be seen the deposition is very poor when a constant ram velocity is utilized. Figure 9 shows the extrusion force time history for the fabricated part with a constant ram velocity. The extrusion force is very inconsistent during the part building process due to system disturbances as mentioned in Part I of this paper such as agglomerate breakdown. The same bar was fabricated in a separate experiment using the same input parameters (e.g., standoff distance, table velocity, horizontal and vertical shifts) except that the extrusion force controller was implemented instead of a constant ram velocity. By comparing the test bars in Figures 7 and 8 it can be seen that the fabrication process is dramatically improved when the extrusion force controller is implemented. Figures 9-11 show the ram velocity and extrusion force time history for the bar fabricated with the extrusion force controller.

Figures 9-11 show the ram velocity and extrusion force time history for the bar fabricated with the extrusion force controller. A reference force of 266 N was chosen due to successful results in previously fabricated parts. The (cid:71) was once again chosen as 0 for a quick reaction to disturbances during the extrusion process. As discussed in Part I of this paper334the nozzle tends to clog if a constant ram velocity is used for an extended period of time. The cause of this is material drying in the die length of the nozzle, thereby reducing the nozzle diameter and eventually causing clogging. By constantly changing the ram velocity the nozzle does not clog. This is believed to be due to variational forcing of the material in the die length of the nozzle. This variation loosens any dried material attached to the nozzle allowing for it to be extruded.

This is believed to be due to variational forcing of the material in the die length of the nozzle. This variation loosens any dried material attached to the nozzle allowing for it to be extruded. This phenomenon is similar to dithering applied to motion systems to overcome the effects of static friction.UnderfillingOverfillingFigure 7: Top view of bars fabricated using extrusion force controller (left) and constant ram velocity (right, v = 2 (cid:80)m/sec).Figure 8: Side view of bars fabricated using extrusion force controller (left) and constant ram velocity (right, v = 2 (cid:80)m/sec).

PART FABRICATION Several parts have been fabricated to demonstrate the feasibility of the developed FEFprocess and the utility of the extrusion force controller. Figures 12–15 show examples of some of the built parts and illustrate the feasibility of the FEF process to make ceramic parts of varying geometries. Figure 12 shows hollow cone and cylindrical geometries. The cones have a sloped wall build angle of 60o. The two unfinished cone pairs show the ability of the FEF process to build geometries with sloped walls without the use of support material. This is possible due to the high solids loading of the ceramic slurry (50 vol.% >). Figure 13 shows two different thin- wall (i.e., single-line) polygonal shapes. There is a buildup of material at the corners due to acceleration/deceleration effects. Tests have been run building thin-wall (single-line) deposited geometries up to a height of 180 mm without the part collapsing.

Tests have been run building thin-wall (single-line) deposited geometries up to a height of 180 mm without the part collapsing. The material has the ability to build taller geometries but, due to limitations of the Z-axis motion, 180 mm is currently the machine’s maximum build height. Figure 14 shows ogive hollow cones after freeze drying. The fabrication of all of these parts was made possible by the utilization of the extrusion force controller, without which good consistent material deposition is not possible.Figure 12: Hollow cones and cylinder (after sintering) fabricated with the extrusion force controller.Figure 13: Thin wall polygonal parts (before binder burnout and sintering) fabricated with the extrusion force controller.33738 mmFigure 14: Hollow ogive cones (after surface finishing) fabricated with the extrusion force controller.5. SUMMARY AND CONCLUSIONSAn on/off feedback controller was designed for improved material deposition.

SUMMARY AND CONCLUSIONSAn on/off feedback controller was designed for improved material deposition. Deposition tests were conducted comparing the deposition consistency and extrusion force with and without the extrusion force controller. Several parts were made with and without the extrusion force controller and the surface finish of these parts was compared.Successful fabrication of parts is not possible without the use of an extrusion force controller. The implementation of the extrusion force controller allows for quick compensation of system disturbances that cause changes in extrusion force. Without the use of feedback control, periodic poor material deposition leads to future deposition problems. Application of a constant ram velocity for material deposition leads to a continuous increase in extrusion force.ACKNOWLEDGEMENTSThis work was supported by the Air Force Research Laboratory under Contract FA8650-04-C-5704.

The authors would also like to thank Mike Hayes of Boeing for his technical support.REFERENCES 1. M. Munson, D. Young, and T. Okiishi, Fundamentals of Fluid Mechanics, Third Edition,New York, John Wiley and Sons Inc., 462-468 (1998).2. B. Russell, J. Lasenby, S. Blackburn, and D. Wilson, “Signal Processing for On-line Control of Paste Extrusion,” AICHE Annual Meeting Conference Proceedings, San Francisco, California, November 16-21, paper 47d (2003).3. B. Russell, “Signal Processing for On-Line Control of Paste Extrusion,” PhD Dissertation, Jesus College, Cambridge, England (2003).4. Y. Li and J. Bridgwater, “Prediction of Extrusion Pressure Using an Artificial Neural Network,” Powder Technology, Vol. 108, 65-73 (2000).338.

eeform Fabrication Assists Forensic Scientists in the Identification of Unknown Victims Bahram Asiabanpour 1α Jerry Melbye β Vicky Melbye β Evan Jensen α Joshua Shaw α α: Ingram School of Engineering β: Forensic Anthropology Center at Texas State (FACTS), Department of AnthropologyTexas State University - San Marcos, Texas 78666Reviewed, accepted September 10, 2008Abstract:According to the International Homicide Investigators Association, there are currently over 40,000 unidentified bodies being held in coroner‟s and medical examiner‟s offices across the United States. Over half are estimated to be victims of homicide, and all are awaiting positive identification. One technique utilized by forensic anthropologists to establish the identity of unknown skeletal or decomposed individuals is the use of facial reproduction. In facial reproduction, soft tissue approximating muscle and skin is added to the skull in an attempt to reflect how the individual looked during life.

In facial reproduction, soft tissue approximating muscle and skin is added to the skull in an attempt to reflect how the individual looked during life. Soft tissue depths at specific locations (i.e., certain craniofacial landmarks) are known and have been standardized according to an individual‟s sex, age, ancestry, and body type. For the most part, facial reproduction is still accomplished manually by adding layers of soft clay to represent tissue on the actual skull of the deceased individual. This method is quite time-consuming (often taking two or more weeks) and often not feasible because the skull may be damaged, or the use of clay overlying the actual skull may destroy evidence. To overcome these limitations, researchers have recently turned to Computed Tomography (CT scan) technology to generate CAD files of unidentified skulls, which are then modeled and recreated with rapid prototyping machines.

One limitation of this method is that it is dependent on the initial CT scanning instrument, which is not portable and requires that the unidentified remains be removed from their original storage sites (typically morgue coolers or crime scenes) to the location where CT scanning instruments are available (often hospital or clinical settings). Because many of these unidentified remains are either skeletonized or are in various stages of decomposition, the transport of these bodies to locations with CT scanning machines is often not possible or permissible. In this paper, we first propose a new method to rapid prototype skulls via stereolithography (STL) files generated by hand-held portable laser scanners, as opposed to using CT scanning machines. These rapid prototypes can then be fabricated for facial reproduction, negating the use of the actual skull, and not requiring the body be removed from its original location.

These rapid prototypes can then be fabricated for facial reproduction, negating the use of the actual skull, and not requiring the body be removed from its original location. Also, results of the facial reproduction for an active case are presented. Secondly, we outline preliminary results of a new computerized facial reproduction and superimposition method, which accurately models tissue depth and is not dependent on the manual application of clay.Keywords: Facial reproduction, Forensic anthropology, skull prototyping, Laser Scanning 1- An introduction to forensic anthropology“Forensic anthropology is the application of the science of physical anthropology to the legalprocess. The identification of skeletal badly decomposed, or unidentified human remains is important for both legal and humanitarian reasons.” [1] Many unidentified body remains are found by law enforcements in a daily basis.

Unidentified bodies may be classified as1Corresponding author, Email: ba13@txstate.eduskeletonized, decomposed, mummified, burned, fresh, living, historic, and nonhuman. Along the US-Mexico border alone, more than 200 unidentified human remains are found annually. The families of these victims may wait for years to hear back about their love ones fate. “Forensic anthropologists apply standard scientific techniques developed in physical anthropology to identify human remains, and to assist in the detection of crime. In addition to assisting in locating and recovering suspicious remains, forensic anthropologists calculate the decedent‟s age, sex, ancestry, stature, pathology and unique features from the skeleton.”[1] Also, forensic anthropologists lend their expertise in the “detection of crimes, and discover evidence of foul play and/or postmortem interval.” [2] Forensic anthropology is very popular among American people and media.

Several popular TV shows such as Bones and CSI have increased the visibility of the forensic anthropology. The drawback is that such semi-fiction programs have increased judges‟ expectation beyond the capacity of this science.2- Facial reproductionAccording to the latest report from the law enforcement agencies, currently more than 40,000unidentified bodies exist in the US [3]. One of the areas that forensic anthropologists aid in the identification of unknown decedents is facial reproduction. In facial reproduction –also known as facial construction, 3D craniofacial reproduction, and facial approximation-, soft tissue is added to skull. Soft tissue depth (i.e., muscle and skin), known as anthropological landmarks, is determined based on the individual‟s race, gender, age, and body type (Figure 1).

Soft tissue depth (i.e., muscle and skin), known as anthropological landmarks, is determined based on the individual‟s race, gender, age, and body type (Figure 1). Currently this method is being done manually by use of adding soft clay as a tissue on the skull.Facial reproduction is one method that is used when the victim has remained unidentifiedthrough the usual avenues to obtain a positive identification. In other words, while facial reproduction has never been accepted in a court of law, it is an investigative tool which can lead to a positive identification. Simply put, the skull is used to generate a facial likeness. This image is then advertised widely, along with information of age, race, sex, stature, and any other unique features found on the skeleton. Positive identification is dependent upon a match between antemortem and postmortem unique characteristics such as fingerprints.

Positive identification is dependent upon a match between antemortem and postmortem unique characteristics such as fingerprints. In the case of decomposed remains, dental x-rays are generally non-existent among these victims. A list of missing persons is then compared against DNA. For example, if a facial reproduction is widely advertised along with the biological profile (age at death, sex, race, stature) this often alerts people to suggest possible missing persons.Figure 1. Anthropological or craniometrical landmarks on a skull (left) and areconstructed facial (right) [4]In current facial reproduction methods, specialists work directly on the skull, however, manytimes, facial reproduction cannot be carried out on a skull because it must be preserved as evidence. By making a prototype of the skull we are not interfering or damaging critical evidence, but are still able to reproduce the face of an unknown victim.

By making a prototype of the skull we are not interfering or damaging critical evidence, but are still able to reproduce the face of an unknown victim. Recently some research has been done to make the facial reproduction in the form of virtual model by use of computer art work (Figure 2) [5 -7].Figure 2. Computer facial reproduction [4]Some experts do not accept computer models as valid reproductions [2]. Another approachthat has recently been proposed is the use of rapid prototyped skull, instead of using the real skull, in the facial reproduction process. The CT image file is then sent to an intermediate program such as Mimics or 3ds Max. This program makes a virtual model of the skull which is then fabricated by a rapid prototype machine. Finally, a facial reproduction specialist builds a face on the prototype by using the clay method [8 and 9]. One limitation of this method is that it is dependent to the CT scan instrument.

Finally, a facial reproduction specialist builds a face on the prototype by using the clay method [8 and 9]. One limitation of this method is that it is dependent to the CT scan instrument. Because many of these unidentified remains are either skeletonized or are in various stages of decomposition, therefore, the transport of these bodies to locations with CT scanning machines is often not possible or permissible.

Also, the CT scanning process is relatively expensive (about $500) and for under budgeted sheriff offices (i.e., US- Mexico border) this expense is not within their budget.3- Facial reproduction on a skull directly prototyped from a laser scanned fileTo overcome the limitations of the previous methods (such as build on the real skull, movingskull from its original site for CT, and also the need for intermediate facial reproduction programs which are not accessible to the public), a method is proposed that simplifies the system while maintaining the accuracy of original skull. The methodology of this approach is as follows: Step 1: 3D scanning of skulls: The skull is scanned by a Polhemus 3D hand-held laser scanner called FastSCAN at the law enforcement agencies or medical examiners‟ offices. This equipment instantly acquires three dimensional surface images when the handheld laser scanning wand is swept over an object, in a manner similar to spray painting.

This equipment instantly acquires three dimensional surface images when the handheld laser scanning wand is swept over an object, in a manner similar to spray painting. FastSCAN works by projecting a fan of laser light on the object while the camera views the laser to record cross-sectional depth profiles. The object's image will immediately appear on the computer screen.Step 2: Post scanning processes and rapid prototyping: The scanned file is exported from the Fastscan software in the stereolithography (STL) format, which is recognizable by the 3D rapid prototyping (RP) machine. The quality of the STL file may be improved or edited within the Fastscan program or intermediate CAD software before sending it to the RP machine. The Rapid prototyping machine used for this process is Zcorp 3D printer model Z450.

The Rapid prototyping machine used for this process is Zcorp 3D printer model Z450. After building a prototype, an exact replica of the skull, it was then reinforced by the Z bond- 101 glue.Step 3: Facial reproduction: Facial reproduction specialist create a clay face from the prototype of the skull by gluing tissue depth markers onto anatomical points of the cast, and clay is used to fill in the spaces between the markers, eyes, nose, lips and hair are added, thus ultimately creating human face. Facial reproduction is not positive identification of an unknown victim; rather, it is a method that can lead to positive identification.Figure 3 illustrates different stages of the proposed method on an active case in San Antonio,TX area. The gap on top of the skull displays areas that suffered blunt force trauma.TraumaFigure 3.

The gap on top of the skull displays areas that suffered blunt force trauma.TraumaFigure 3. Laser scanning, CAD file improvement, rapid prototyping and different stages of the facial reproduction in an active case involving a White or Hispanic female 18-24 years old4- Advantages and limitationsThe advantages and uniqueness of this approach is that the 3D laser scanning can be doneon-site and does not require an intermediate/confidential program for building a virtual model. After a positive identification is obtained, the prototyped skulls may be used as teaching tools in forensic anthropology labs across the United States. .The drawback for this method is that some common laser scanning processes and rapid prototyping machines have limitations. For example, the quality of darker surfaces is not great and in the end we may see gaps in certain areas of the prototype.

For example, the quality of darker surfaces is not great and in the end we may see gaps in certain areas of the prototype. Because of the fixed position of the skull during the scanning process, some areas of skulls such as under lower jaw, neck, and inside of the mouth, are not reachable by the hand-held laser scanner gun. The software for the Zcorp 450 rapid prototyping machine, Zprint, does not give much flexibility to the user to define the thickness of the model to be prototyped. Therefore, the entire skull is built as a solid material, rendering it very heavy and costly to produce.5- Supporting toolsTo address some of the above limitations, some tooling additions and process modifications have been implemented. A new mechanism has been designed, manufactured, and tested. This mechanism holds the skull at minimum contact points and is capable of turning the skull and the laser scanner transmitter simultaneously in all six turning directions (+ and – around X, Y, and Zaxes).

The advantage of this tool is that it allows the operator to scan the entire surface from all sides. (Figure 4).Figure 4. Skull holding/turning mechanism for the laser scanning processTo improve the quality of the scanned skull CAD file (e.g., filling the gaps between surfaces)and to build the prototype with the desirable thickness, Materialise 3matics software was used. This enabled us to save considerable amount of material and build a much lighter skull. This model later was built by use of Zcorp Z450 rapid prototyping machine (Figure 5).Figure 5. New skull prototype with half inch Zcorp 450 machineThe use of the interfacial programs such as 3Matics has facilitated the skull prototyping in better surface quality. However, the visual inspection of the final prototype hypothesizes that surface quality improvement process might have changed the skull geometry and orientation of the important landmark points.

However, the visual inspection of the final prototype hypothesizes that surface quality improvement process might have changed the skull geometry and orientation of the important landmark points. As an alternative method, a new project is under investigation by the authors and a team of talented high school students participating at Texas State University‟s Siemens honor summer mathematics camp. In this research, the research team is trying to give desirable internal thickness to the skull shell that was produced by the laser scanning process by creating a freeform geometry inside the skull. The result of this output will be a file in STL format acceptable for rapid prototyping.

The result of this output will be a file in STL format acceptable for rapid prototyping. Since the external shell of the skull is untouched, the geometry of the external surface of the reproduced skull will be exactly similar to the original skull.6- ValidationThe newly opened Forensic Research Facility at Texas State University is the largestdecomposition facility in the world, and one of three facilities in operation [10]. At this facility donated bodies will be laid out in the open area to decompose naturally. This facility providesvery unique opportunity for validation of this research. The human body, facial features, and head can be laser scanned and measured precisely before being placed in the open to decompose. Once the skull has skeletonized, it can be laser scanned again. A prototype can be produced from the CAD for facial reproduction.

Once the skull has skeletonized, it can be laser scanned again. A prototype can be produced from the CAD for facial reproduction. The final reproduction can then be compared against a photo of the living individual and measured visually, manually, or through software work such as the three dimensional quantification facial shapes by Thomas [11, Chapter 4].7- Photographic superimpositionPhotographic superimposition is the process of overlaying two photographs, one of the unidentified skull, and one of a possible victim. This method is usually used when investigators believe that the remains may belong to the person in the photograph [12]. If there is a match, the person is considered to be a possible victim; however, if there is no match, the person is excluded.A preliminary effort has been done to implement a more consistent photographic superimposition.

Based on this research, by use of parametric CAD software a picture was imaged and matched with a designed (not scanned) face (Figure 6). This match was successfully built by the Zcorp 450 color 3D printer (Figure 7).Figure 6. Photographic superimpositionFigure 7. A physical prototype of a model built by photographic superimposition8- Computerized facial reproductionSeveral research and developmental works have been reported on computerized facialreproduction systems. In a method by Tu et. al., a given skull is matched with a face that is saved in a database of CT scanned heads [13]. Also, law enforcement in the United Kingdom and theFBI are using computer programs for facial reproduction called 3ds Max c and Reface c respectively [8]. Clement and Marks in their 2005 book, “Computer-Graphic Facial Reproduction” have presented different facial reproduction approaches [11].

Clement and Marks in their 2005 book, “Computer-Graphic Facial Reproduction” have presented different facial reproduction approaches [11]. Subsol and Quatrehomme use Crest lines extracted from a CT scan file, registration of the feature lines, geometric normalization, and computing the 3D transformation [11, Chapter 5]. Also, Vargas and Sucar propose a method for predicting the most probable facial features by use of discrete landmarks, Bayesian networks, mathematical morphology, and computer graphics [11, Chapter 7]. Davy et.al. have introduced a facial reproduction method using the 3ds Max computer modeling software. They use anatomical reproduction and Non-Uniform Rational B-Splines (NURBS) Control Vertex (CV) curves based on the underlying muscles in modeling (Neave method) [11, Chapter 10].Under a collaborative work between the School of Engineering and the Department of Anthropology at Texas State University, a new facial reproduction method is being developed.

This method, similar to Davy et. al., is based on the underlying muscular or Neave method. Also, the facial reproduction will be fully automated in generating, orienting, adjusting landmarks, muscles, and skin based on information such sex, age, race, and body type of the victim with minimum artist and/or forensic anthropologist interaction.The proposed method for the facial tissue regeneration program includes the following steps.First, all known tissue thickness data will be stored in a database with the ability to query based on the race, gender, age, and build of an individual. Then, a computer program will be used to create generic facial tissue geometry, which will also be added to the database. In the initial stages of development, the user will interact with the software to pick points best suited to match the landmarks for tissue data thickness.

In the initial stages of development, the user will interact with the software to pick points best suited to match the landmarks for tissue data thickness. The points will then be processed to compute a normal vector for each point in order to provide the proper location and orientation for the tissue generation process (Figure 8). The generic tissue will then be attached to the points in their corresponding locations, and will be automatically oriented to fit the imported skull model. Then, through a set of graphical menus, tissue thickness data are modified by the desirable combination of race, sex, age, and build. As the user inputs the data, the tissue thickness data will be updated on the 3D model, and start to resemble the owner of the unidentified skull. The user will be able to pick from a set of predefined features such as ears, noses, hair, and eyes.

The user will be able to pick from a set of predefined features such as ears, noses, hair, and eyes. An algorithm will then be run to wrap skin around the facial features, and apply race specific skin texture and eye color for a near photorealistic rendering of the subject. Once complete, the model can be prototyped in full color, or rotated in 3D space for help with identification.Figure 8. Vectors to specify anthropological landmarks‟ locations and orientations9- Conclusions and discussionAccording to Stephan and Henneberg, approximately1 in 16 reconstructed faces have led to the identification of a missing person or unknown victim [14]. The traditional method of manual facial reproduction is quite time-consuming and often destroys evidence (the skull). The use of CT scan technology to generate CAD files of unidentified skulls, which are then modeled and recreated with rapid prototyping machines, has improved the facial reproduction process.

The use of CT scan technology to generate CAD files of unidentified skulls, which are then modeled and recreated with rapid prototyping machines, has improved the facial reproduction process. However, this method is dependent on the non-portable CT scanning instrument and necessity of removing the unidentified remains from their original storage sites to the CT scanning instrument, which is often not possible or permissible. In this paper, a new method to rapid prototype skulls via STL files generated by hand-held portable laser scanners is proposed. These rapid prototypes can then be fabricated for facial reproduction, eliminating to the need to move the unidentified remains from their original storage sites.References:1- American Board of Forensic Anthropology website, Retrieved on March 28, 2008http://www.csuchico.edu/anth/ABFA/2- Komar, D.

EEFORM FABRICATION OF 3D ZINC-AIR BATTERIES AND FUNCTIONAL ELECTROMECHANICAL ASSEMBLIESEvan Malone, Kian Rasa, Daniel Cohen, Todd Isaacson, Hilary Lashley, Hod LipsonComputational Synthesis Laboratory, Department of Mechanical and Aerospace Engineering, Cornell University, Ithaca NY 14853 USAem224@cornell.edu Reviewed, accepted August 19, 2003AbstractThis paper reports on a fabrication platform and extensions to deposition-based processes that permit freeform fabrication of three-dimensional functional assemblies with embedded conductive wiring and power sources. Structure and joints are produced by fused deposition of thermoplastics and deposition of elastomers. Conductive wiring is achieved by deposition of various low-melting-point alloys and conductive pastes. Batteries based on zinc-air chemistry are produced by deposition of zinc, electrolyte, and catalysts, with separator media and electrodes.

Batteries based on zinc-air chemistry are produced by deposition of zinc, electrolyte, and catalysts, with separator media and electrodes. Details of the deposition processes are provided and several printed assemblies are demonstrated.IntroductionIn the last several years, there have been an increasing number of solid-freeform fabrication processes and materials being researched. It has not escaped the notice of some that this diversity is nearing the point at which almost any conventionally manufactured part could be made entirely via SFF processes, and that a synthesis of compatible processes may permit the fabrication of entire functional assemblies [Weiss and Prinz, 1998; Safari et. al., 2000] and even complete functional systems [Lipson et. al., 2000]. The Cornell Computational Synthesis integrated, Laboratory (CCSL) electromechanical systems, and an understanding of the interface between the design process and this new fabrication domain.

The Cornell Computational Synthesis integrated, Laboratory (CCSL) electromechanical systems, and an understanding of the interface between the design process and this new fabrication domain. A high-performance motion control platform and an initial set of tools have been designed and constructed that have enabled the exploration of layered fabrication processes in a wide variety of materials and material combinations. The materials experimented with thus far include thermoplastics, low-melting-point alloys, and a variety of gels and slurries. Here we report on freeform fabrication of a zinc-air battery which is capable of powering a small DC brush motor, thermoplastic and elastomer flexure joints, two approaches to embedded wiring, and structures with embedded wiring.

These components are some of the basic building blocks necessary for freeform fabrication of three-dimensional, functional, electromechanical systems.isinvestigating the freeform fabrication of functional,Fabrication PlatformIn order to provide maximum freedom for experimentation with a wide variety of materials and processes, a custom robotic platform and two material deposition tools have been designed and constructed. The initial requirement employed in the design of these tools can be stated briefly as follows: the motion control platform should provide maximal parametric freedom to the deposition processes, and permit the sequential use of many deposition processes in the course of fabricating a given object. The most significant assumption informing the design is that materials will be deposited as streams or droplets in a layered manufacturing process.

The most significant assumption informing the design is that materials will be deposited as streams or droplets in a layered manufacturing process. Positioning is therefore limited to three Cartesian axes, and an emphasis was placed on velocity363regulation, path-following and positioning accuracy, resolution, repeatability, and high acceleration to achieve fine features while printing at a constant material feed rate. Lower acceleration performance would require more complex deposition feedrate control. Table 1 describes the performance requirements employed in the design of the positioning system.Figure 1: Gantry robot fabrication platformA Cartesian gantry-configuration robot (Figure 1) has been selected for this application for its relatively simple control, large payload capacity relative to rigidity, and positioning/path- following performance.

Linear motors, ballscrews, and cable-pulley systems were all considered for the drivetrain, and ballscrews were selected for simplicity. A cable-pulley system is an attractive alternative because the drive motors themselves do not need to be located on the moving components of the system.Table 1: Fabrication System Performance SpecificationsMinimum material stream/drop diameter Materials cross-section area Build rate Nominal speed along path Min turn radius at nominal speed Tool Position Accuracy (+/-) Tool Position Repeatability (+/-) Positioning Resolution Build envelope x Build envelope y Build envelope z Max XY acceleration250 µm 4.9 x10-8 m2 2.5 x10-9 m3/s 0.05 m/s 125 µm 25 µm 25 µm 5 µm0.3 m 0.3 m 0.3 m 20.75 m/s20.010in0.55in^3/h4.92 x10-3 in 9.84 x10-4 in 9.84 x10-4 in 1.97 x10-4 in 11.8 in 11.8 in 11.8 in 2.12 gA software application has been created to manage path planning and control.

Multi- material objects are defined using multiple STL (stereolithography) files, each describing a single material. Each material is associated with a deposition tool and material properties governing layer thickness, deposition width and deposition rates. Geometry slicing and path generation algorithms construct unique perimeter contour and fill raster paths based on the tools’ and materials’ parameters, and combine the layers into a fabrication sequence with increasing364height. Special care is needed to prioritize layers of similar heights according to interaction among materials.Deposition ToolsA freeform fabrication system capable of producing complex, functional products requires a highly versatile deposition tool or set of deposition tools to address the need to deposit a broad range of materials with vastly different physical properties.

The set of materials explored in the freeform fabrication of batteries can be organized into four basic categories: plastics, metals, liquid chemicals, and chemical pastes. For the initial experiments, two separate deposition tools were designed – one for plastics and metals, and another for liquid chemicals and pastes. The tools were designed to be compatible with both a Cartesian robot and a six-axis robotic arm in order to allow for experimentation with both types of positioning systems.

The tools were designed to be compatible with both a Cartesian robot and a six-axis robotic arm in order to allow for experimentation with both types of positioning systems. Load restrictions and dimensional limitations of robotic arms require that tools be small and lightweight, a feature that also leads to lower systemic inertia and enhanced robotic performance.(a)(b)Figure 2: Extrusion tools constructed: (a) wire-fed extruder, and (b) syringe toolThe deposition tool for metals and plastics (Figure 2a) is an extruder that feeds material in wire form (0.050” – 0.070” diameter) through an actively air-cooled metal guide tube and into a heated liquefier block containing a nozzle [Swanson et. al. 1999]. The tool has been successfully tested with solid core Pb-Sn solder wire, as well as ABS (acrylonitrile-butadiene-styrene) thermoplastic wire.A separate tool was designed to deposit liquids and chemical pastes (Figure 2b).

This tool accepts standard commercial 10cc Luer-lock syringe barrels and plungers. The plunger position is actuated by a linear stepper-motor capable of exerting 50 lbf. Prior experiments with a pneumatic dispensing system proved open-loop pressure control to be unsuitable for freeform deposition of slurries; instead, volumetric control provided by the current approach is more broadly applicable, and also allows for more precise reverse-flow control. The design of the tool simplifies experimentation with a broad range of materials. Materials may be easily changed by substituting a different syringe barrel and plunger in the plunger-driver component of the tool.365Freeform Fabrication of a Zinc-Air BatteryThe freeform fabrication of a functional battery was selected as a comprehensive test of the multiple material capabilities of the previously described platform.

There is little literature on the production of freeform, three-dimensional energy storage devices, though a planar thin- film cell has recently been demonstrated by Power Paper Ltd. (2003) for low-power applications. Freeform fabricated batteries are a major step toward freeform fabricated, active, functional, electromechanical systems. The size, shape, and performance of a cell become free design parameters, and can be customized for the specific geometry and functional requirements of a product. A cell can be placed at arbitrary locations within the product, rather than forcing the design of a product to be limited by the characteristics of commercially available energy sources. With this in mind, research and experiments were conducted in an effort to produce a freeform cell.

With this in mind, research and experiments were conducted in an effort to produce a freeform cell. For the purpose of these experiments, a cell is defined as a device that converts the chemical potential energy between its anode and cathode materials into electrical energy by means of redox reactions: reduction (electron gain) at the cathode, oxidation (electron loss) at the anode. A battery is comprised of one or more connected cells.

The essential components of a Zn-air cell are the negative terminal, anode, separator, cathode catalyst, cathode, positive terminal, and electrolyte (Table 2, Figure 6).Table 2: Key Battery Components / Materials Tested*Cell ComponentMaterials TestedElectrolyte0.25 Molar -> 8 Molar solution of potassium hydroxide (KOH) and distilled waterNegative terminalPaste of methylcellulose (MC) with copper (Cu, 99% purity 2-5 µm), or with silver (Ag, 99% purity 1 µm)AnodeSlurry of electrolyte with zinc (Zn, 97.1% purity, dust) and surfactantSeparatorPaper, or Rescor 740 insulating ceramic foam (Cotronics Inc.)Cathode catalystSlurry of carbon black, manganese dioxide (MnO2, 80-85% purity), and electrolyteCathodeAirPositive terminalPaste of methylcellulose with nickel (Ni, 99% purity -325 mesh), or copper, or silverMany of these materials are hazardous, and should only be handled with proper training and protective equipmentZinc-air cell chemistry was the first considered because of its simplicity and high energy density.

The basic chemical reactions are:At Air Cathode: ½ O2 + H2O + 2e (cid:314)2OH- At Zn Anode: Zn (cid:314) Zn2+ + 2eZn2+ + 2OH- (cid:314) Zn(OH)2 Zn(OH)2 (cid:314) ZnO + H2O______________________________________ Overall:Zn + ½O2 (cid:314) ZnO : E0=1.65 VThe experiments outlined below investigate the effects on cell performance of electrolyte concentration, current collector composition and geometry, cell structure and construction, and cathode catalyst composition. After the initial tests were conducted and the fundamental characteristics understood, the research shifted to adapting a battery design to solid-freeform fabrication processes.366Zn-Air Electrolyte ExperimentsTests were conducted with the aim of understanding the basic anode and electrolyte chemical ratios and concentrations, and all other variables were held constant.

The test procedure [Isaacson, 2003] included mixing the Zn with various concentrations of KOH (0.25M to 8M) and placing the mixtures in an ABS cell casing, using copper wire as both the negative and positive terminals (Figure 3a).(a)(b)Figure 3: Experimental cells, in cross-section: (a) basic Zn-air cell and (b) Zn-air cell with MnO2 and carbon black catalystPaper was initially used as a separator for simplicity and consistency. The cells were loaded with a 102 Ohm resistor and the voltage recorded. It was found that an electrolyte concentration of 7M to 8M gave the cell a more consistent voltage over its lifetime.Catalyst ExperimentsManganese dioxide and carbon black were then added to the cell to improve cell performance [Tinker, 2002]. The ratio of MnO2 to carbon was derived from manganese dioxide alkaline cells. These materials are also used as a catalyst in commercial Zn-air batteries [Linden et. al. 2002; Duracell, 2001].

The ratio of MnO2 to carbon was derived from manganese dioxide alkaline cells. These materials are also used as a catalyst in commercial Zn-air batteries [Linden et. al. 2002; Duracell, 2001]. By weight, the anode was comprised of 65% Zn and 35% 8M KOH, the catalyst was comprised of 50% MnO2, 44% 8M KOH, and 6% carbon black, and the separator layer consisted of 8M KOH saturated paper. The cell was constructed in a similar way to the previous cell design with the addition of a catalyst layer (Figure 3b), and tested in the same manner. Important observations made from these tests were that the consistency of cell performance improved and the voltage doubled compared to previous tests. The most significant observation was that the power output increased by 500%.

The most significant observation was that the power output increased by 500%. To reduce the potential energy barrier at the terminal junctions, nickel and copper wire were used for the cathode and anode current collectors respectively.Surface Area and Current Collector ExperimentsTests were conducted to understand the effects of surface area and increase overall performance. Using the same cell chemistry and mass of material as in previous tests, the cell surface area was increased from a 1” diameter to a 3” diameter, and the current collectors/terminals were doubled in length (6” coiled wire to a 12” coiled wire). The cells were loaded with a 33 Ohm resistor to acquire data relevant to the goal of running a 30mW electric DC brush motor. The graphs (Figure 4) below demonstrate that cells of this simple design are367able to supply more than 30mW at 1V for more than ½ hour.

The graphs (Figure 4) below demonstrate that cells of this simple design are367able to supply more than 30mW at 1V for more than ½ hour. The cells were able to turn the motor, and in doing so, validated the tested cell chemistry and structure.Voltage Voltage vs. Time: Increased Surface Area & Current CollectorPowerPower vs. Time: Increased Surface Area & Current Collector1.20.0451.150.040.035s t lo V1.11.0510.95) s t t a W(r e w o P0.030.0250.020.0150.010.90.0050.8501251 501 751 1001 1251 1501 1751 2001 2251 2501 2751 3001 3251 35011254 507 760 1013 1266 1519 1772 2025 2278 2531 2784 3037 3290 3543Time (seconds)Time (seconds)(a)(b)Figure 4: (a) Voltage vs. time of cell with increased surface area and current collector area and (b) power vs.

time of cell with increased surface area and current collector area and (b) power vs. time of cell with increased surface area and current collector areaSeparator ExperimentsMethylcellulose (MC) gel and Rescor 740 insulating ceramic foam were both tested for their efficacy as separator layers, to replace the only remaining unprintable component in the cells – paper. Methylcellulose (3 grams of MC to 100mL of 100°C distilled water) does not interact well with 8M KOH electrolyte [Danko, 2003] - the gel deteriorates and allows shorting between cathode and anode. Rescor 740 (100 parts base to 64 parts activator) is a permeable ceramic foam which absorbs the electrolyte and allows for the migration of ions across the layer, making it a viable choice as separator.Extrusion Experiments Conductive PastesVarious metallic pastes and solder alloy were tested as materials for current collectors.

Pb-Sn solder proved to be incompatible with the cell chemistry and was ruled out as a terminal material. MC was mixed in 1:1 ratios by mass with Cu, Ni, and Ag powders, and each was extruded through a syringe creating extrudable wires. Silver paste proved to be the best material for the cell terminals for two reasons; the silver terminals did not react with the cell chemistry and had lowest electrical resistivity of the pastes. This paste was thus used for both the anode and cathode current collectors/terminals. The nozzles used for testing were 14Ga stainless steel needles, which produced a stream approximately 1.4mm in diameter.SeparatorSurfactant was added to the Rescor 740 ceramic foam material to enable extrusion and to delay the curing of the material while in the syringe. The foam slurry is highly viscous and prone to phase separation, and requires a large diameter nozzle for successful extrusion.

The foam slurry is highly viscous and prone to phase separation, and requires a large diameter nozzle for successful extrusion. Initial testing indicated that the addition of the surfactant in the separator material reduces power output slightly, and more the optimum material and/or Rescor/surfactant ratio. The internal diameter of the nozzle used was 1.9mm, and the resulting stream of material was about 1.3mm in diameter.testingis necessaryto determine368AnodeThe anode material (a slurry of Zn and KOH solution) was exceptionally difficult to extrude, as Zn settles out of the slurry and clogs the syringe. Surfactant was added to prevent clogging, but the resultant increase in ease of handling comes at the expense of some cell power output. Therefore, the slightest possible amount of surfactant was used. The anode consisted of 60% Zn, 30% KOH, and 10% surfactant by weight.

Therefore, the slightest possible amount of surfactant was used. The anode consisted of 60% Zn, 30% KOH, and 10% surfactant by weight. The nozzle diameter was approximately 7mm, resulting in comparably-sized drops of deposited material.CatalystThe MnO2 and carbon black powders are extrudable when pre-mixed with KOH solution. In order to achieve an homogeneous slurry, the catalyst material was forced through a syringe several times before use, using a pneumatic dispensing machine. 14Ga stainless steel syringe needles were used, and these produce a stream of about 1.4mm in diameter.Freeform Fabricated Cell DesignThe next round of experiments focused on the fabrication process. These tests were designed to demonstrate that 100% of the materials for a complete cell could be extruded, to verify that freeform fabricated cell performance would be adequate (i.e.

These tests were designed to demonstrate that 100% of the materials for a complete cell could be extruded, to verify that freeform fabricated cell performance would be adequate (i.e. sufficiently powerful to turn the motor and have comparable voltage/power output to previously tested cells), and to identify problems that might occur during production with the fabrication platform.Table 3: Process for Constructing a Manually Extruded CellNegative TerminalThe silver slurry was extruded into a container with a copper wire exposed to the silver for data collection purposes. It was then dried using a heat gun. The heat helps to evaporate the water from the MC solution and allows the material to be strong enough to receive the load of the next layer.AnodeHere the Zn-KOH and surfactant slurry was added over the silver negative terminal.SeparatorFor testing purposes, the Rescor separator layer was pre-cast.

It was allowed to cure at room temperature for approximately five hours to ensure that the layer was solid and would not interfere with the other cell components and chemistry.Catalyst and Cathode TerminalFinally, after the separator was placed over the Zn, the catalyst was extruded freely over the separator. The silver slurry was then liberally extruded over the catalyst and separator. A copper wire (not shown) was placed over the silver cathode terminal for data collection purposes.369Manually Extruded Cell and ResultsFirstly, a cell was made by hand using only the extrudable materials and extrusion deposition. This process is detailed in Table 3 above. The graphs below (Figure 5) show the voltage and power output of the cell - loaded with a 100 Ohm resistor - over a one hour period. The all-extruded cells have satisfactory performance, delivering more than 10mW at more than 1V for the entire hour of data collection.

The all-extruded cells have satisfactory performance, delivering more than 10mW at more than 1V for the entire hour of data collection. The downward spike in the graphs occurred when the cell was connected to the motor for demonstration testing. Multiple tests conducted with this cell design and manufacturing technique show repeatable results over a sample size of approximately six cells, suggesting the feasibility of producing a functional cell on the freeform fabrication platform.Voltage vs. Time: Manually Extruded CellPower vs.

Time: Manually Extruded CellPower vs. Time: Manually Extruded Cell1.60.020.0181.40.0161.20.01410.012s t l o V0.8s t l o V0.010.0080.60.0060.40.0040.20.00201149 297 445 593 741 889 1037 1185 1333 1481 1629 1777 1925 2073 2221 2369 2517 2665 2813 2961 3109 3257 3405 35531147 293 439 585 731 877 1023 1169 1315 1461 1607 1753 1899 2045 2191 2337 2483 2629 2775 2921 3067 3213 3359 3505Time (seconds)Time (seconds)(a)(b)Figure 5: Performance of manually extruded cell; (a) voltage, and (b) power to 100 Ohm load.Freeform Fabricated CellThe freeform fabricated cell design is essentially identical to that of the manually fabricated test cells. Figure 6 shows a cross-section view of the cell design.

Figure 6 shows a cross-section view of the cell design. The zinc anode is surrounded by silver paste to enclose the zinc and prevent KOH evaporation, and to increase the reaction surface area of the negative terminal.Figure 6: Cross-sectional view of freeform fabricated cell designThe separator layer is intentionally designed to be excessively large compared to the rest of the cell components to ensure that the cell does not short. The catalyst layer is extruded on top of the separator layer, and the cathode terminal, constructed of MC/silver paste, is printed directly370over the catalyst. Both the anode and the cathode terminals are connected to the motor or the data acquisition board via copper wires.

Both the anode and the cathode terminals are connected to the motor or the data acquisition board via copper wires. Figure 7 highlights some key steps in the successful freeform fabrication of this cell design.(a)(b)(c)(d) Figure 7: Sequence of operations in freeform fabrication of battery: (a) solid view of battery model in CCSL SCM software, (b) slicing of battery model in SCM application to generate tool paths, (c) deposition of Rescor separator layer, (d) close-up photograph of completed battery, and (e) dissection of battery (anode terminal on left and separator layer/zinc anode on right)(e)Results and DiscussionThe performance benchmark test for the freeform-fabricated cell was the ability to power a small DC brush motor, which consumes about 30mW, unloaded. The cell was able to provide the starting current for the motor, and ran the motor for about 2 seconds. Figure 8 depicts the open circuit voltage of the cell after it ran the motor.

The cell was able to provide the starting current for the motor, and ran the motor for about 2 seconds. Figure 8 depicts the open circuit voltage of the cell after it ran the motor. When comparing this voltage to previous tests it is evident that the cell has excellent construction (no internal shorting) because the open circuit voltage is near the theoretical maximum potential for a cell of this chemistry (~1.6V). After having run the motor, however, the cell was no longer able to generate significant power, perhaps because of poor oxygen transport into the cell, electrolyte evaporation, or insufficient zinc quantity. A thinner separator layer, enclosing the cell in a thermoplastic case, and increasing the volume of deposited zinc anode will be examined as remedies for each of these problems, respectively. It is also apparent that the dimensions of the physical cell differ from those specified in the design.

It is also apparent that the dimensions of the physical cell differ from those specified in the design. This is a result of the number of materials involved, material preparation variability, material property degradation, and high sensitivity of tool calibration to material properties. Further refinements in equipment calibration and material processing methods will improve output control.371Voltage Open Circuit Voltage vs. Time: Freeform Fabricated Cell1.81.61.41.2s t l o V10.80.60.40.20150099914981997249629953494Time (seconds)1 in.Figure 8: Open circuit voltage of freeform fabricated batteryFigure 9: ABS and RTV silicone elastomer flexure jointOther Material/Functionality ExperimentsAs a demonstration of a multi-material functional mechanical assembly, a flexure joint was fabricated using ABS as the rigid end members and a 1-part, room-temperature vulcanizing (RTV) silicone as the flexible connection (Figure 9).

The silicone is filled with carbon black to make it freestanding upon extrusion. The combination of electrical and mechanical functionality within the same freeform fabricated part is highly desirable and a number of approaches have been investigated [Ting et. al. 2001; Safari et. al. 2000]. Some initial qualitative experiments with applications of MC / metal powder pastes as functional materials outside of batteries have revealed that dehydration of the MC gel can lead to shrinkage and cracking of a deposited road of MC paste, and that adhesion to substrates also suffers after dehydration. A rudimentary electromechanical assembly was freeform fabricated as lines of silver / MC paste embedded in an ABS and silicone flexure joint (Figure 10c).

A rudimentary electromechanical assembly was freeform fabricated as lines of silver / MC paste embedded in an ABS and silicone flexure joint (Figure 10c). This device successfully carried sufficient current to light an LED (~10mA), but was too delicate to survive much mechanical use due to cracking and detachment of the conductive paste.Solder alloys are being investigated as a means of depositing wiring into components, and fabricating metal parts [Priest et. al., 1997], and at least one commercial process exists which is capable of depositing solder alloy wiring onto a wide variety of substrates [Hayes et. al., 1998]. Our previous experiments revealed that a large reservoir of molten eutectic alloy, for instance in a syringe, has the tendency to drain uncontrollably from the reservoir or to freeze in a nozzle.

Our previous experiments revealed that a large reservoir of molten eutectic alloy, for instance in a syringe, has the tendency to drain uncontrollably from the reservoir or to freeze in a nozzle. As an alternative approach, a solid-core, Pb-Sn solder wire was used as the feedstock for the wire-fed extrusion tool, and solder deposits were made directly on a build surface. Figure 10a depicts the results of these tests. It was found that there are small, separate regions of the parameter space in which it is possible to form either lines of overlapping frozen droplets of approximately 1mm in diameter (Figure 10a, left), or a very thin but continuous wire of approximately 250µm in diameter (Figure 10a, right). As a test of compatibility between materials, several ABS and solder test coupons were freeform fabricated (Figure 10b).

As a test of compatibility between materials, several ABS and solder test coupons were freeform fabricated (Figure 10b). The deposited solder is electrically continuous for the entire “U” shape in the right-hand sample which was produced at a fast feed rate, but only continuous about 2/3 of the total length of the “U” for the left-hand sample, which was produced at a much slower solder feed rate. The lead- tin solder does not wet ABS well, so making a robust interface between the two materials372requires some surface preparation, or simply tightly embedding the solder within a channel in the ABS material.

There is some indication that wetting may be improved by the use of indium- bearing alloys, although this has yet to be tested.(b)(a)(c)Figure 10: (a) Effect of deposition parameters on quality of solder alloy wiring, (b) partial success at embedding solder alloy wiring in thermoplastic part, and (c) flexure joint with silver/methylcellulose paste conductorConclusionsA zinc air cell has been successfully manufactured via freeform deposition techniques. The printed cell delivered at least 30mW for 2sec, from only 1gm of zinc slurry. Mechanical flexure joints have been freeform fabricated, as have a thermoplastic part with embedded metal wiring and a flexure joint with embedded silver/methylcellulose paste wiring which is capable of carrying more than 10mA of current. Obvious routes exist to improving the functionality and durability of all of these.

Obvious routes exist to improving the functionality and durability of all of these. The battery, flexure joints, and embedded wiring together demonstrate the feasibility of combining multiple materials into fully functional assemblies within a single freeform fabrication platform. Through further testing and optimization, higher performance cells are achievable, as are more sophisticated functional assemblies, leading the way to the production of immediately useable, freeform fabricated, electromechanical devices.AcknowledgmentsThis work was supported in parts by the U.S. Department of Energy, grant DE-FG02- 01ER45902. Thanks to Douglas Charles of Inco Technical Services for information on mixing ratios of methylcellulose and metal particles, and to Cordelia Rasa for information about methylcellulose products. D.C.

Thanks to Douglas Charles of Inco Technical Services for information on mixing ratios of methylcellulose and metal particles, and to Cordelia Rasa for information about methylcellulose products. D.C. thanks the NASA/NY space grant program for its support.373References Danko, Thomas, Strength Properties of Separators in Alkaline Solutions, Viskase Corporation, Chicago,2003Duracell Incorporated, 2001, “Duracell zinc Air Batteries, Material Safety Data Sheet.”Fuller, S., Willhelm, E., Jacobson, J., “Ink-Jet Printed Nanoparticle Microelectromechanical Systems,”Journal of Microelectromechanical Systems, 11, no. 1, (Feb 2002): 54-60Harris, Frank, and Sieminski, Dennis, “Recent Advances Enable Primary zinc-Air for Portable Devices”Wireless Symposium Spring 1999 Conference, San Jose Ca, February 22-26 1999Hayes, D.J., Cox, W.R., Grove, M.E., “Micro-jet printing of polymers and solder for electronicsmanufacturing,” J. Electron. Manuf. (Singapore), 8, no. 3-4, (Sept.-Dec.

L. Bourell, J. J. Beaman, H. L. Marcus, R. H. Crawford, and J. W. Barlow, eds.), The University of Texas at Austin, August 2001Weiss, L. and Prinz, F., 1998, "Novel Applications and Implementations of Shape DepositionManufacturing," Naval Research Reviews, Office of Naval Research, Three/1998, Vol. L374.

eeform Fabrication of a Complete Electromechanical RelayEvan Malone, Hod LipsonMechanical and Aerospace Engineering, Cornell University evan.malone@cornell.edu Reviewed, accepted September 5, 2007AbstractResearch PaperPurpose We seek to produce electromechanical relays entirely via freeform fabrication as part of a larger effort to freeform fabricate complete electromechanical devices. Relays which can switch using an input current of less than 100 microamperes, current gain greater than 10, and an open/closed resistance ratio of greater than 103 will make feasible the control of freeform fabricated actuators by printable organic polymer transistor circuits, opening up a design space of completely freeform fabricated electromechanical actuation systems. Design/methodology/approach We conceived a simple relay design which makes use of our prior successes in freeform fabrication of ionomeric polymer- metal composite actuators (IPMC).

Design/methodology/approach We conceived a simple relay design which makes use of our prior successes in freeform fabrication of ionomeric polymer- metal composite actuators (IPMC). The electrically conductive surface electrodes of an IPMC serve to selectively close or open a load circuit as the actuator deforms, and the mold used for casting the actuator materials also serves as the relay housing. We are seeking candidate switch contact materials, IPMC electroding materials, candidate housing and contact designs, and manufacturing sequences to find combinations which result in functional IPMCs which when hydrated release themselves from the load contacts while maintaining signal contact, yet have low contact resistance with the load contacts. Candidate housings with contacts were fabricated using a Fab@Home Model 1 desktop fabricator.

Candidate housings with contacts were fabricated using a Fab@Home Model 1 desktop fabricator. We manually cast variants of IPMC materials into these housings, and tested these partially freeform fabricated relays using a PC-based data acquisition system. Findings Partially freeform-fabricated relays achieved a controllable load path resistance change of a factor of two at switching frequencies up to 0.05 Hz with a 1.2V control input and an RMS input current of 3.8mA. At 1.2V load potential, load/input current gain is roughly 1.05. The device functions above 1Hz, but current gain falls below 1. Research limitations/implications Incompatibilities between the tested relay contact materials and the IPMC materials lead to IPMCs of low quality which have internal shorting, resulting in reduced force output, and poor switching performance. Our “wet-contact” design limits devices to operation below 1.2V for both the control signal and the load. Device gain is barely greater than 1.

Our “wet-contact” design limits devices to operation below 1.2V for both the control signal and the load. Device gain is barely greater than 1. Additional design and materials optimization is required. Practical implications This work demonstrates the feasibility of completely freeform fabricating electromechanical relays, and suggests that the stated performance goals are attainable with further work. Originality This work constitutes the first demonstration of complete, functional relays which can feasibly be produced entirely by freeform fabrication without manual assembly.Keywords electromechanical, relay, IPMC, biomimetic, robotics, freeform fabrication, rapid prototyping, printableIntroduction The frontier of research in freeform fabrication is pushing outward to encompass ever larger and smaller dimensions, ever more materials, and now increasingly toward fabricating in multiple materials with multiple processes, all within the same fabrication system.

We claim that the implicit and valuable drive behind all of this research is to reduce the cost and complexity of infrastructure and the level of technical training required to build essentially anything – both to expand the realm of what can be built, and to allow smaller companies, even individuals, to build or invent almost anything for themselves. The ultimate expression of this drive is a compactmanufacturing system including materials and software tools which together will make the building of a functional device as simple and accessible as printing a digital photograph is today.We have been working toward this goal both from the direction of simplifying and popularizing multimaterial SFF technology, and through developing hardware, software, and materials which allow the integral freeform fabrication of complete electromechanical devices, food, engineered tissues – essentially anything.

Along the first line, we have created the Fab@Home Project(Malone and Lipson, 2007), which develops and publishes open-source, open-architecture, multimaterial SFF system designs – including hardware, software, materials, and applications ideas - and promotes them to the public via a user-modifiable wiki-website, collaborative open- source software development, online discussion groups, and support of commercialization efforts. Pursuant to the second avenue of research, we have demonstrated freeform fabrication of batteries (Berry et al., 2005), structures, joints and wiring (Malone et al., 2004), soft polymer “artificial muscle” actuators (Malone and Lipson, 2006a), and living tissue constructs (Cohen et al., 2006).

The burgeoning fields of organic electronics and “printed” electronics have demonstrated that complex integrated circuits can be made entirely through solution processes (Nilsson et al., 2005, Paul et al., 2003), and from our preliminary efforts at producing single transistors with our fabrication systems (Malone and Lipson, 2004), we believe it will soon be possible for us to fabricate integrated circuits with them as well.Figure 1. The "electronics gap": freeform fabricated actuators have input current and voltage requirements compatible with a freeform fabricated battery, but exceed the current capacity of organic transistors by 3 orders of magnitudeOur own frontier of research is now moving beyond demonstrating the fabrication of single device types toward multi-device integration – for example, an entirely freeform fabricated actuator subsystem including a battery, IPMC actuator, and organic electronics control circuitry.

While the batteries we have been able to fabricate can supply sufficient current to actuate an IPMC, that current is three orders of magnitude larger than what can be controlled by a printable transistor. This “electronics gap” (Figure 1) would require us to build 1000 transistors in parallel to control one actuator, or to find or develop some kind of amplifier which can be completely freeform fabricated.We have conceived of an electromechanical relay which employs an IPMC as the mechanical switch, and which in principle can fill our electronics gap. Here we present the relay design, and partial freeform fabrication and characterization of some units.BackgroundStandard electromechanical relays are binary amplifiers – a small electrical current supplied to a coil electromagnetically actuates a mechanical armature as a switch, which in turn controls a much larger current through some useful load.

The ratio of the change in load current from “on” to “off” to the current required to actuate the relay can be defined as the relay’s (current) gain, G. Maximum load current is limited by resistive power dissipation in the metal-to- metal switch contacts, while minimum coil current is limited by the need to move the switch armature quickly to minimize switching noise and/or arcing at the load contacts – modern DC relays have G~100 (Omron, 2006, Tyco-Axicom, 2006).Relays have been produced before in part through SFF processes (Palmer et al., 2006), but that work involved metal machining processes as well as assembly of parts.

Here, we seek to produce such a device entirely through additive manufacturing processes that are compatible with all of the other freeform fabricated devices we can produce so that a single compact fabrication system can build up complete electromechanical devices without any assembly of parts.An Ionomeric Polymer-Metal Composite (IPMC) actuator is a branched polymer- electrolyte (ionomer) film which has been given parallel electrically conductive surface electrodes. When saturated with an appropriate solvent, an electric field applied via these conductive surface electrodes causes the IPMC to bend toward the more positive electrode through a combination of electrostatic (mutual repulsion of anion-terminated branches) and ion- driven pressure gradient (electrophoresis of solvated mobile cations) mechanisms (Nemat-Nasser and Li, 2002, Shahinpoor and Kim, 2001).

After this initial rapid motion, and depending upon the choice of solvent (Hyoung-Tae et al., 2006, Nemat-Nasser and Zamani, 2003), ionomer (Nemat-Nasser and Li, 2002), and possible nanocomposite additives(Nguyen et al., 2006, Paquette et al., 2003), the IPMC may or may not relax back toward the negative electrode, putatively because of solvent back-diffusion (Nemat-Nasser and Zamani, 2003).The standard approach to IPMC fabrication (Kim and Shahinpoor, 2003), and that which results in the highest actuator performance, involves purchasing a solid membrane from a commercial source, and applying surface electrodes to it via metal-salt reduction to metal nanoparticles, usually followed by a chemical or physical metal plating to reduce surface resistance.As first demonstrated by Kim and Shahinpoor (Kim and Shahinpoor, 2002), it is also possible to produce IPMCs entirely through solution processes.

Cast IPMCs can be made to greater thickness than commercially available membranes which allows an increase in force output (Kim and Shahinpoor, 2002, Nguyen and Yoo, 2007), although at least anecdotally, solution processing of the ionomer seems to result in reduced electromechanical efficiency relative to IMPCs made with commercial (extruded) membranes (Kim and Shahinpoor, 2001). Based on these successes, we developed formulations of Electroding or Active Materials (formulations with “EM”, and “AM” designations) and solution processable ionomer Interior Materials (formulations with “IM” designations), all of which are syringe dispensable and tolerate being cast into silicone molds, and succeeded in producing both molds and IPMCs in a single freeform fabrication process (Malone and Lipson, 2006a).

This is significant because itallows an IPMC to be integrally fabricated with other devices, such as batteries and wiring, as part of an electromechanical subsystem or complete electromechanical device. One material particularly important for this success is an electrically conductive elastic vapor barrier material, designated EM19, which serves as an electroding material.

One material particularly important for this success is an electrically conductive elastic vapor barrier material, designated EM19, which serves as an electroding material. EM19 not give our IPMCs the longest reported operation in air without rehydration (> 4 hours, 3000+ actuation cycles (Malone and Lipson, 2006a)) by preventing solvent loss through the actuator surfaces, but also improves the quality of films of the solution processed ionomer material when cast into silicone molds by acting as a hydrophilic intermediary.Given that an IPMC generally has high-conductivity electrode surfaces, and can be caused to bend by a relatively small electrical signal, it can act as both coil and armature in a relay if it can be actuated in order to selectively close a load circuit with sufficiently small resistance (Figure 2).

Our concept is to fabricate a relay housing which contains electrical contacts for the load circuit and for the IPMC and has a recess into which the IPMC materials can be cast and the entire housing annealed to remove the water solubility of the materials. Additional contacts can then be deposited atop, the cavity filled with water and capped, creating a sealed device (to prolong the IPMC service life) with three or more electrical terminals.Figure 2.

(a) Schematic of IPMC relay concept; (b) “wet contact” design in which load contacts are within the solvent-filled housing; (c) “dry-contact” design in which IPMC is extended with an external conductor which passes through the housing to interact with the load contacts in dry conditions.Several design issues must be considered which affect the performance and mode of operation of any relay, including whether it should be normally conducting or not (normallyclosed or normally open), how to maximize the current gain (minimize the input current requirement, minimize the load current when the device is not conducting, and minimize the resistance to the load current when the device is conducting), and how to achieve rapid switching. Several additional issues must be considered when the relay will employ an IPMC as the switching armature, and the entire device must be additively manufactured with no assembly.

Several additional issues must be considered when the relay will employ an IPMC as the switching armature, and the entire device must be additively manufactured with no assembly. These include: whether the load contacts should be wet or dry (as the IPMC must be wet to operate), whether an IPMC of good quality, with good dimensional tolerances and high performance can be cast into the same structure that will house the relay, and whether the IPMC, once cast from solution, can free itself from its mold automatically so that it can operate without needing manual removal and reassembly.A very simple design for a “normally closed” IPMC relay would be one in which one IPMC electrode (the “anode”) is grounded and acts as the load path to ground (Figure 2(a)).

If in the unactuated “relaxed” state, the IPMC closes the load circuit, then a positive potential applied to the upper electrode (the “cathode”) will bend the IPMC upward, breaking the load circuit.A “normally open” version cannot be made simply by placing the load contact above the upper electrode, as in this configuration the signal current source and load current source must be the same, and G > 1 is not possible. Maximizing gain is a matter of ensuring that the IPMC is of good quality – hence high electromechanical efficiency, and that the contact resistance between the load contacts and the armature (IPMC anode, or perhaps some conductor attached to the IPMC) is as small as possible when conducting, and as large as possible when not. This is largely a material properties issue (e.g. intrinsic resistivity, hydrophobicity if the contacts are wet), but also depends upon the pressure that the IPMC can apply to close the load circuit.

This is largely a material properties issue (e.g. intrinsic resistivity, hydrophobicity if the contacts are wet), but also depends upon the pressure that the IPMC can apply to close the load circuit. Switching speed is tied to manufacturing tolerances (how far the IPMC must move to open/close the circuit), electromechanical efficiency of the IPMC (force/charge transfer), and the allowable control signal amplitude (in water, limited to < 1.24V, above which electrolysis occurs, wasting power and corroding electrodes; other solvents can tolerate higher potential). Wet contacts are conceptually simpler (Figure 2(b)) – one of the IPMC electrodes (which will be wet) acts to close the load circuit by pressing against the load contacts, but many materials have high contact resistance when wet, and wet contacts limit the load supply voltage to < 1.24V for electrolysis reasons.

In addition, when one of the IPMC electrodes acts as part of the load’s path to ground, the load current will generate a resistive potential in this electrode, possibly counteract the input control signal to the IPMC. Dry contacts (Figure 2(c)) require encapsulating the IPMC to prevent solvent loss (which sacrifices efficiency) and/or coupling some external conductor to the IPMC to switch the load circuit – complicating the manufacturing process. The quality of cast IPMCs, and whether they self-release from the mold/housing is dependent upon the IPMC material formulations, additives, etc., and the choice of mold/housing materials.ExperimentalIn our prior work on freeform fabrication of IPMCs, we employed a large, non-portable multimaterial freeform fabrication system (Malone et al., 2004).

One drawback of that system was that at the time we lacked fume extraction facility proximate to the system, and hence had to transport IPMC specimens to a fume hood and back to the machine repeatedly during the fabrication process, complicating the positional registration of fabrication steps. For the freeform fabrication processes in the present research, we are employing a Fab@Home Model 1 personal fabrication system (Malone and Lipson, 2007) which is compact and portable enough to setup ina wetlab environment. This permits us to extract fumes directly from the system without removing the samples. The Model 1 dispenses materials from disposable 10cc polyethylene syringe barrels (EFD, Inc.) by the action of a linear stepper motor coupled to a disposable neoprene syringe piston (EFD, Inc.).

The Model 1 dispenses materials from disposable 10cc polyethylene syringe barrels (EFD, Inc.) by the action of a linear stepper motor coupled to a disposable neoprene syringe piston (EFD, Inc.). The software application accepts 3D model data, allows the user to associate each model component with a material type, and then plans and executes the multimaterial fabrication process based on user supplied material data, prompting the user to change the material at the appropriate points in the process. Material changes are simply a matter of removing one syringe barrel and replacing with another.

Material changes are simply a matter of removing one syringe barrel and replacing with another. More details on the Model 1 can be found online (Malone et al., 2007).Micrographs of IPMC specimens were captured with an AmScope ZM-1TZ-FOR-GT-3M Stereo Inspection Microscope with 3-megapixel camera, (AmScope, Inc.).The structural material employed for IPMC molds and relay housings is GE Silicone II RTV (General Electric, Inc.) The candidate anode, cathode and load contact materials included Leit-C CCC, a commercial conductive carbon cement (SPI Supplies, Inc.), SPI 05002-AB, a high purity silver paint (SPI Supplies, Inc.), and SS26F, a silver-filled silicone RTV (Silicone Solutions, Inc.).

Candidate IPMC electrode materials (designation EM or AM) included AM1, carbon black-filled Nafion composite (Malone and Lipson, 2006b), EM19, a thermoplastic rubber/silver colloid composite material (Malone and Lipson, 2006b), EM22, Ag-nanoparticle-filled Nafion composite, and EM23, a composite of 0.7-1.3µm Ag powder-filled Nafion. All candidate IPMC internal materials are based on a 20wt% “liquid Nafion” dispersion in a blend of alcohols and water (Sigma-Aldrich #527122), which has been converted to Li+ cation-form via a LiOH treatment (Malone and Lipson, 2006b).

IM1 (Malone and Lipson, 2006b), , with 50wt% N,N- Dimethylformamide (Sigma-Aldrich) as a film-quality additive, and 2 wt% of fumed silica, and IM2, which is 50wt% of the 20Li+ .The details of materials formulation and methods for freeform fabrication of IPMCs can be found elsewhere (Malone and Lipson, 2006b).We performed automatic testing of complete relay devices using a PC-based data acquisition system (DAQBoard 2000 PCI, IOTech, Inc.) and DASYLab software (measX GmbH & Co. KG). Data is collected at 100Hz sampling rate, and data analysis performed with MATLAB (The MathWorks, Inc.).Figure 3 provides a schematic of the automatic test setup. The relay load path (e.g. from load terminal 1 through the IPMC anode – if it is in contact with the load terminals – to load terminal 6) is connected as the lower leg of a voltage divider, with the resistive load, RLoad=115Ω +/-1%.

We use separate, isolated, power supplies for the load and the IPMC actuation signal to ensure that the load circuit resistance signal is not contaminated by current leakage from the IPMC control signal. All testing is performed with VLoad = 1.2VDC, and an IPMC input square wave signal with 1.2V amplitude and frequency ranging from 0.005Hz to 1Hz.Figure 3.

All testing is performed with VLoad = 1.2VDC, and an IPMC input square wave signal with 1.2V amplitude and frequency ranging from 0.005Hz to 1Hz.Figure 3. Data acquisition setup for automatic testing of IPMC relay; RLoad=115Ω, VLoad = 1.2VDC at all times, and the IPMC input signal is a square wave of 1.2V amplitude and frequency typically in the range 0.005-1 Hz.Candidate Materials SelectionGiven the aforementioned relay design issues of contact resistance, self-release, and IPMC performance, we evaluated candidate materials for the relay housing, anode, cathode and load contacts, and IPMC electrodes.We performed a simple evaluation of relative contact resistance between IPMC electrode materials and candidate contact materials by pressing cast specimens of IPMC electrode materials against two electrically isolated traces of candidate contact material deposited manually on a glass slide.

A small mass (2.2g) is used to generate a consistent, moderate pressure between the materials for a more fair comparison, and the total path resistance from one trace, through the electrode material, and back to the other trace was measured with a digital multimeter.Measurements were performed with the materials fully dry, and again with them saturated in deionized water, in order to inform the tradeoff between wet and dry contact relay designs. Figure 4 shows the data from these experiments.Leit C CCCElectrode Material to Contact Material Path Resistance0.11101001000100001000001000000EM22 DryEM22 WetEM19 Dry EM19 WetAM1 DryAM1 WetElectrode MaterialContact Path Resistance (Ohm)SS26FSPI 05002-ABFigure 4.

Data from wet and dry resistance tests of candidate IPMC electrode materials against candidate relay load contact materialsNote that the data are not adjusted to compensate for the resistance of the traces themselves or for the instrument contact to the traces, and were used for relative comparisons of material pairs only. It is clear that in dry conditions, either the EM22 or the EM19 electrode material against the SPI 05002-AB silver paint is the best combination, while for wet conditions, the combination of EM19 and SS26F is superior. The Ag-nanoparticle filled electrode material, EM22, is not included in these tests as it was found to be essentially non-conducting, and hence not a viable material for closing the load circuit.Figure 5.

Poor casting results for IPMC materials in (Design 1) relay housings constructed of GE Silicone II RTV, and electrical contact material SS26F; note the thick rims, curled and/or cracked surfacesWe examined the issues of self-release and IPMC quality by proceeding to fabricate candidate relay housing designs, and then casting IPMC materials into them. We evaluate self-release qualitatively by observing whether the IPMC is in substantial contact with the walls of the housing, and quantitatively by measuring the resistance from the anode terminal to each of the load terminals. IPMC quality is evaluated qualitatively by visual inspection for cracks, pinholes, and visible anode to cathode shorting and by optical microscopic inspection of internal structure, and quantitatively by measuring the anode to cathode resistance.(a)(b)Figure 6.

Optical micrographs of a poor-quality cast IPMC extracted from relay housing; (a) top down view showing irregular surface and impressions left by load contacts, scale bar 1mm; (b) edge-on view of cleaved IPMC showing well-formed and separated layers of EM19, EM23, and IM1 toward center (left), but anode to cathode shorting at edge (right), scale bar 500µm.Design 1 (Figure 5) uses the single load contact design shown in Figure 2(a), and the majority of the housing base surface area consists of SS26F – roughly ½ of the area for the anode contact and ½ for the load contact. With this design, it became apparent fairly quickly that the IPMC materials do not cast well atop the SS26F/GE Silicone II combination of which the housings are constructed (Figure 5).

With this design, it became apparent fairly quickly that the IPMC materials do not cast well atop the SS26F/GE Silicone II combination of which the housings are constructed (Figure 5). None of the previously viable IPMC material combinations resulted in high-quality IMPCs – all tended to have very thick and cracked rims which were frequently the sites of anode to cathode shorting, and irregular internal material layer thickness (Figure 6).

Several design modifications were used to attempt to improve the casting quality and to reduce cathode-to-anode shorting, including reducing the fraction of the surface area occupied by the SS26F material (Design 2), switching to the multiple load-contact design (Design 3) shown schematically in Figure 2(b), adjusting the height of the conductive contacts relative to the inside bottom surface of the housing/mold (Design 4), and depositing the anode electrode layer prior to depositing the housing walls, so that the anode layer could perfectly cover the bottom surface of the housing cavity, but the cavity walls are completely devoid of anode material (Design 4 variant).After each design modification, a new set of set of relay housings was produced with the Fab@Home Model 1. We cast IMPC materials by pipette into these housings, varying the choices of electrode material(s), internal materials, and the number of layers (from 3 to 5).

We cast IMPC materials by pipette into these housings, varying the choices of electrode material(s), internal materials, and the number of layers (from 3 to 5). All casting attempts were dried at 35C in flowing air, and then annealed at 70C for 45 minutes. After annealing, all devices were inspected to select those with the best IPMC quality. These were made into complete relay devices by hand, by manually depositing (from syringe) a cathode insulator (of GE Silicone II) and a cathode terminal (of SS26F or EM19). For all finished devices, the resistance between anode and cathode, and between anode and electrode was measured with the device at ambient humidity, and also fully saturated with deionized water.Finished devices have 36AWG copper wire bonded to their anode and cathode terminals with silver paint (SPI 05002-AB) to improve the repeatability of electrical connection to the data acquisition system.(a)(b)(c)Figure 7.

Design 4 - successful IPMC-based relay; (a) an exploded view of Design 4, showing the sequence of fabrication from bottom to top; (b) assembled view with cap and IPMC removed to show multiple load contacts and mold-recess in housing; (c) four examples of Design 4 relays with automatically fabricated housings and contacts, manually cast IPMCs, manually deposited cathode contact and contact insulator, and bonded electrode wires for testing. Note: the lid is not included in these test samples to allow for post-testing disassembly and analysis.In total, to date, we have performed 27 different casting attempts of 12 different IPMC material/stacking options, in 4 different housing designs. Of the 27, 8 were deemed to be of good enough quality for completion to testable relays. In testing these, we have thus far had only 1 definite success, with marginally detected response in a few other instances. Figure 7 shows the most successful design of housing.

In testing these, we have thus far had only 1 definite success, with marginally detected response in a few other instances. Figure 7 shows the most successful design of housing. It includes 6 total load terminals to increase the probabilityof finding a pair that experience the largest change in resistance as the actuator oscillates during testing, and the heights of the terminals have been designed protrude about 500µm above the interior of the housing cavity floor as a compromise between excessive disruption of the planarity of IPMC materials during casting, and insufficient protrusion to make contact with the IPMC electrode. The cathode contact insulator, which is deposited immediately prior to the cathode contact, prevents the cathode contact material from flowing around the edges of the IPMC and shorting to the anode contact.The most successful single device employed the combination: housing, Design 4; 5- layers (EM19, EM22, IM1, EM22, EM19); cathode conductor material, SS26F.

Results of testing this device are presented below.Results / Discussion1502000.01 Elapsed Time [s]Load Path Resistance [Ohm]Load Path Resistance vs. Time0.02501000Figure 8. Successful test of partially freeform fabricated relay; IMPC input signal 1.2V, 0.05Hz square wave, VLoad = 1.2VDCFigure 8 shows time-series data of current input to the IMPC, and load path resistance calculated from the load circuit voltage divider potential. The input signal is a 0.05Hz, 1.2V square wave. The RMS input current to the IMPC, IRMS,In, for this particular sample is 0.0038A, and the RMS power, PRMS,In, is 0.0039W. The open/closed resistance ratio achieved is roughly 2 (~300Ω/150Ω). With a minimum resistance, Rmin of 150Ω, the current gain, G, for this device is 1.05. The large “glitch” at t ~ 110s is unexplained.2070904010030500.01600.005 Elapsed Time [s]Load Path Resistance [Ohm]Load Path Resistance vs. Time8010Figure 9.

The large “glitch” at t ~ 110s is unexplained.2070904010030500.01600.005 Elapsed Time [s]Load Path Resistance [Ohm]Load Path Resistance vs. Time8010Figure 9. Successful test of partially freeform fabricated relay; IMPC input signal 1.2V, 1Hz square wave, VLoad = 1.2VDCAs can be seen in Figure 9, at 1Hz, this same device has a smaller response. IRMS,In = 0.0031A, PRMS,In = 0.0035W. The resistance ratio becomes 1.2 (~220Ω/180Ω), and G ~ 0.39.Future Work The present results are not sufficient to fill our need for an amplifier which can connect the microampere domain of organic polymer semiconductors to the milliampere domain of the IPMC used as a macroscale actuator. Given the difficulties we have experienced with casting of IMPC materials into the housings we have fabricated, we propose to explore additional viable housing and electrical contact materials, including epoxies and hydrophilic thermoplastic polymers both with and without conductive fillers.

Improvements to IPMC casting will improve the dimensional tolerances and electromechanical efficiency of the resulting IPMCs, enabling a reduction of their size and input power requirements and thus the current gain of the relay. IPMC performance can be enhanced through nanocompositing with activated silica (Nguyen et al., 2006, Paquette et al., 2003), which will also reduce the relay input current, as well as the driven actuator’s input requirements.Conclusion We are confronted by significant barrier to our goal of developing multimaterial freeform fabrication systems capable of “printing” complete electromechanical devices.

We can print batteries capable of supplying sufficient power for the actuators we can print, and it is feasible to produce integrated circuits entirely through solution processing within our fabrication system.However, the foreseeable circuits, based on organic polymer semiconductors, will not be able to directly control the level of power required to operate a meaningful actuator. Hence there is a gap between what needs to be controlled and what can do the controlling. To bridge this “electronics gap” we are developing a relay which employs an IPMC actuator in lieu of the electromagnet and armature of a normal DC relay. Here we have demonstrated complete, functional IPMC-based relays produced entirely by solution/extrusion processes, with the relay housing and electrical contacts fabricated automatically with a Fab@Home Model 1 system, and the remainder deposited by hand from pipette or syringe.

Despite substantial effort at optimizing materials and exploring designs, our current material suite does not yet yield the level of performance that we require, but we are actively pursuing several obvious approaches to enhancing device performance, and confident of eventual success. Relays with sufficiently high performance will allow us to fabricate complete actively controlled electromechanical devices within a single, compact freeform fabrication system, without assembly, opening up an entirely new design space for engineering and product innovation.Acknowledgments This work was supported in parts by the U.S. Department of Energy, grant DE-FG02- 01ER45902.References BERRY, M., MALONE, E. & LIPSON, H. (2005) Freeform Fabrication of Zinc-Air Batteries with TailoredGeometry and Performance. IN BEAMAN, J. & BOURELL, J. (Eds.) 16th Solid Freeform Fabrication Symposium. Austin, TX, USA.COHEN, D. L., MALONE, E., LIPSON, H. & BONASSAR, L. J.

eeform Fabrication of Aluminum Alloy Prototypes Using Laser MeltingHideki KYOGOKU1, Masashi HAGIWARA2, and Toshifumi SHINNO1 1 Faculty of Engineering, Kinki University Higashihiroshima, Hiroshima 739-2116 Japan 2 Aspect Inc. Inagi, Tokyo 206-0802 JapanAbstractIn this study, a direct selective laser sintering/melting machine was designed and constructed. The machine has a 50 W Yb-fiber laser, a galvanometer scanner and a powder delivery and build system. It was confirmed that the machine works well. The fabrication conditions of aluminum alloys were investigated using the machine. The optimum laser power, scan speed and scan pitch were investigated by experiments. The effect of addition of metal powder as additives on laser scanning process was investigated to fabricate the sound laser-scanned body of aluminum alloys based on Al-12Si alloy.

The effect of addition of metal powder as additives on laser scanning process was investigated to fabricate the sound laser-scanned body of aluminum alloys based on Al-12Si alloy. It was found that the smooth single-scan track can be fabricated at lower laser power and higher scan speed by the addition of a laser absorption material. An aluminum alloy prototype was successfully produced using optimum laser scanning conditions.IntroductionSelective laser sintering/melting has been widely utilized as one of the RP processes of various materials such as polymers, metals and ceramics. In the field of automotive industries especially, the RP process has been introduced to develop a new car as fast as possible. Also it is necessary for recent new cars to be designed and manufactured with light weight to save energy. Therefore, the application of aluminum alloys has increased in automotive parts.

Also it is necessary for recent new cars to be designed and manufactured with light weight to save energy. Therefore, the application of aluminum alloys has increased in automotive parts. The RP technology, however, is not able to manage well to produce trial products of aluminum alloy. The authors developed a laser sintering/melting machine which can fabricate metal products, in particular aluminum alloy and titanium alloy products, with high density and high accuracy, against the background of development of high performance fiber-laser and alloy powders in the project. Simultaneously, some companies developed commercialized machines. The features of these machines are using a fiber-laser to fabricate high accuracy products and having a chamber to keep a vacuum atmosphere.Recently, research on the RP process of aluminum alloys has been reported. Sercombe et al. [1-3] reported on the fabrication of a resin-bonded aluminum alloy part using selective laser sintering.

Sercombe et al. [1-3] reported on the fabrication of a resin-bonded aluminum alloy part using selective laser sintering. But sound parts of aluminum alloy were not fabricated by this method because the sintering performs in the atmosphere of nitrogen gas. Yu et al. [4] reported the infiltration condition of aluminum alloy parts. Taminger et al. [5] and Mahale et al. [6] investigated the fabrication conditions of dense aluminum parts by electron beam melting and fabricated 2024 and 7075 aluminum alloy sample parts. The research on the fabrication of aluminum alloy parts using direct selective laser melting or laser sintering, however, has not been reported.In this study, a laser sintering/melting machine was designed and constructed to produce aluminum alloy parts. The fabrication conditions of aluminum alloys were investigated using the constructed machine. The fabrication conditions, such as laser power, scan speed, scan pitch and layer thickness, were investigated by experiments.

The fabrication conditions, such as laser power, scan speed, scan pitch and layer thickness, were investigated by experiments. Finally, an automatic transmission part was fabricated in the optimum condition.Development of Laser Sintering/Melting MachineA laser sintering/melting machine was designed and constructed to produce aluminum alloy parts. The machine consisted of a 50 W Yb-fiber laser (wave length:1090 nm, spot size: around 170µm), a galvanometer-scanner and a computer-controlled powder delivery and build system. The laser and galvanometer-scanner were computer-controlled using the STL data converted from the CAD product data. The inside of the machine is shown in Fig. 1. The powder delivery and build system has a recoater, two powder supply unit and a laser sintering unit. The system can coat a 50 µm layer of powder (minimum). The chosen atmosphere can be changed to argon gas.

The system can coat a 50 µm layer of powder (minimum). The chosen atmosphere can be changed to argon gas. The performance of this machine was confirmed.Fig.1 Laser sintering/melting machine developed in the projectExperimental MethodPowder characteristicsIn this study, gas-atomized Al-12Si powder and a laser absorption material were prepared. The mean particle diameter of the Al-12Si powder was around 33µm. The Al-12Si powder had irregular shape, as shown in Fig.2. These powders were mixed using a V-blender or a ball mill.

The Al-12Si powder had irregular shape, as shown in Fig.2. These powders were mixed using a V-blender or a ball mill. The addition of laser absorption material was 0.1~ 4 mass%.30 μmFig.2 Shape of Al-12Si powerExperimental procedureThe mixed powders were filled into the container of the powder supply unit and thenlaser-scanned with the Yb-fiber laser under various fabrication conditions as follows:Laser power: 10~50WScan-speed: 5~20 mm/s - Scan-pitch: 0.1~0.3 mm - Layer thickness: 0.1~0.4 mmThe surface morphology and microstructure of the laser-scanned bodies were examinedby an optical microscope and a scanning electron microscope with EDX.Results and DiscussionsFabrication conditions (1) Effect of laser absorption material and the change in morphology of single-track with laser power and scan speedIn order to produce sound parts, it is very important to examine fabrication conditions by the constructed machine.

Especially, laser power, scan-speed, scan-pitch and layer thickness are significant factors [7-10]. In this study, the optimum conditions of these factors were examined experimentally and then the single-track process map was produced.At first, the optimum conditions of the laser power and the scan-speed were examined to obtain a smoothly continuous single-track. In the case of no laser absorption material, almost unmelted tracks or partially melted single-tracks were obtained. Thus, it was found that a smoothly continuous single-track could not be fabricated using only Al-12Si powder. Therefore, a laser absorption material was added.Figure 3 shows the change in morphology of single-tracks with laser power and scan-speed in the case of the addition of 0.1% laser absorption material. As shown in Fig.3, in the case of 0.1% addition, a partially melted single-track was obtained in almost conditions.

As shown in Fig.3, in the case of 0.1% addition, a partially melted single-track was obtained in almost conditions. It was, however, found that a smoothly continuous single-track can be obtained by the addition of more than 0.5% in consideration of the relation between laser power and scan-speed. The width of single-track of aluminum alloy is fairly narrow compared to that of other metals such as alloy steels and is close to the laser spot diameter. This may be because only the laser-scanned part was melted due to high thermal conductivity of aluminum alloy. As an example of a smoothly continuous single-track, the process map of the addition of 2.0% is shown in Fig.4.

As an example of a smoothly continuous single-track, the process map of the addition of 2.0% is shown in Fig.4. Here, the shape of a single-track is roughly classified into three types as follows:(a) a smoothly continuous track (b) a melted and balled continuous track (c) a partially melted trackFigure 4 also shows the change in morphology of single-tracks with laser power and scan speed in the case of the addition of 2.0% laser absorption material. Comparing to the case of 0.1% addition, the area in which a smoothly continuous single-track can be obtained widely expands in the process map as ahown in Fig. 4(a). It is found from this process map that a smoothly continuous single-track can be obtained at a laser power of more than 30 W and at a scan-speed of less than 15 mm/s. The relation between laser power and scan-speed for smoothly continuous single-tracks changes with the addition of laser absorption material.

The relation between laser power and scan-speed for smoothly continuous single-tracks changes with the addition of laser absorption material. The area in which a smoothly continuous single-track can be obtained became narrower at an addition of more than 3%.Thus, laser power and scan-speed which smoothly continuous single-tracks can beobtained are found easily by the process map.Also, the effect of energy density, E, calculated by the following equation wasexamined.Jmm2Ｅ=ＰｖｄWhere P=laser power [W], v=scan-speed [mm/s] and d=laser spot size [mm]. The values calculated as 0.17 mm in laser spot size are shown in Fig.4(a).

The values calculated as 0.17 mm in laser spot size are shown in Fig.4(a). It is found from thisfigure that it is difficult to obtain a smoothly continuous single-track at an energy density of less than approximately 10 J/mm2.(a) Relation between laser power and scan speed▲ partially melted track600µm200µm600µm200µm(10 W, 10 mm/s, 5.9 J/mm2) (40 W, 10 mm/s, 22.5 J/mm2) (b) Change in shape of single-scan tracks with scan speed Fig.3 Change in morphology of single-track with laser power and scan speed (0.1% laser absorption material)(2) Fabrication conditions of the surfaceThe change in surface state of the laser-scanned body with scan-pitch was examined. The scan-pitch was examined between 0.1～0.3 mm. The scanning pattern is shown in Fig.5. The fabrication of the smooth surface was difficult with a scan-pitch of 0.2 mm and 0.3 mm because the laser spot size is around 0.17 mm.

The scanning pattern is shown in Fig.5. The fabrication of the smooth surface was difficult with a scan-pitch of 0.2 mm and 0.3 mm because the laser spot size is around 0.17 mm. As mentioned above, this is because the width of single-track of aluminum alloy is close to the diameter of laser spot. As a result, it was found that the optimum scan-pitch is between 0.1mm and 0.13 mm. This value is approximately 1/3 overlap of the width of track which is similar to that of other metals.The laser-scanned surface state is affected not only by scan-pitch but also addition of laser absorption material, laser power and scan-speed. In the case of the addition of 4.0% laser absorption material, the laser-scanned body was cracked due to large shrinkage.

In the case of the addition of 4.0% laser absorption material, the laser-scanned body was cracked due to large shrinkage. Therefore, the addition content of laser absorption material was set at 2%.(a) Relation between laser power and scan speed■ Continuous track ● melted and balled track600µm200µm600µm(40W, 5mm/s, 47.1 J/mm2) (20W, 5mm/s, 23.5 J/mm2)▲ partially melted track600µm200µm(10W, 10mm/s, 5.9 J/mm2) (b) Change in shape of single-scan tracks with scan speed Fig.4 Change in morphology of single-track with laser power and scan speed (2.0% laser absorption material)Fig.5 Laser scanning pattern (mm)200µmFigure 6 shows the change in the surface state of the laser-scanned body with laser power at a scan-speed of 20 mm/s and at a scan-pitch of 0.116 mm. As shown in this figure, the smooth surface of the laser-scanned body can be fabricated at a laser power between 30 W and 50 W. The surface of the body laser-scanned at 40 W is most smooth with fewer pores.

The surface of the body laser-scanned at 40 W is most smooth with fewer pores. Figure 7 shows the change in the surface state of the body laser-scanned with a scan pitch at a laser power of 40 W and at a scan-pitch of 0.1 or 0.116 mm. As shown in this figure, the smooth surface of the laser-scanned body can be fabricated at a scan-speed between 15 and 25 mm/s.As a result, it was found that the smooth surface of the laser-scanned body can be fabricated at a laser power between 30 W and 50 W, at a scan-speed between 15 mm/s and 25 mm/s and at a scan-pitch between 0.1 mm and 0.13 mm.600µm200µm600µm200µm(a) 30 W (b) 40 W600µm200µm(c) 50 W Fig.6 Change in surface state of the laser-melted body with laser power (laser absorption material:2%, scan-speed:20 mm/s, scan-pitch:0.116 m)Aluminum alloy prototype production (1) Effect of layer thicknessIt is important to examine the effect of layer thickness to fabricate a sound body.

In this research, the effect of layer thickness was examined between 0.1 and 0.4 mm. In the case of test specimens, a sound body could be obtained at a layer thickness of 0.25mm. As mentioned above, the optimum fabrication conditions are as follows:Content of the addition: 2% - Laser power: 40 W - Scan-speed: 20 mm/s - Scan-pitch: 0.116 mm -An automatic transmission part in 1/3 size was fabricated under these conditions as shown in Fig.8. As shown in Fig.8(b), delamination took place. This may be because of weakness of interlaminar strength against thermal deformation. Therefore, in order to improve the interlaminar strength, the layer thickness decreased to 0.18 mm which is nearly equal to the laser spot diameter.

Therefore, in order to improve the interlaminar strength, the layer thickness decreased to 0.18 mm which is nearly equal to the laser spot diameter. As a result, as shown in Fig.8(c), the layered body did not delaminate, but cracked.600µm200µm600µm(a) 15 mm/s (b) 20 mm/s600µm200µm(c) 25 mm/s Fig.7 Change in surface state of laser-sintered body with scan speed (laser power:40 W, scan-pitch:0.1 or 0.116 mm)(a) CAD data (b) Layer thickness 0.25 mm(c) Layer thickness 0.18 mm (d) Layer thickness 0.18 mm (two times scanning) Fig.8 An aluminum alloy prototype of automatic transmission part200µm(2) Effect of laser annealingAs mentioned above, cracks occurred with increasing thickness of the layered body even in the body fabricated in the optimum condition. Therefore, laser annealing was performed after first laser scanning. As shown in Fig.8(d), an aluminum alloy product without crack could be fabricated by applying laser annealing.

Therefore, laser annealing was performed after first laser scanning. As shown in Fig.8(d), an aluminum alloy product without crack could be fabricated by applying laser annealing. Thus, it was found that laser annealing is so effective to fabricate a sound body of aluminum alloy prototype.ConclusionsIn this study, a direct selective laser sintering/melting machine was designed andconstructed to produce aluminum alloy parts. (1) The performance of the designed and constructed machine was confirmed. (2) It was found that a smoothly continuous single-track can be obtained by the addition of more than 0.5% of absorption material. (3) Laser power and scan-speed with smooth continuous single-tracks can be obtained when using the process map. (4) It was found that the smooth surface of the laser-scanned body can be fabricated at a laser power between 30 W and 50 W, at a scan-speed between 15 mm/s and 25 mm/s and at a scan-pitch between 0.1 mm and 0.13 mm.

(5) It was found that choice of layer thickness which is nearly equal to the laser spot diameter and the application of laser annealing are effective to fabricate a sound layered body of aluminum alloy.Thus, an aluminum alloy prototype was successfully produced using optimum laserscanning conditions.AcknowledgementsThe project was performed by the following members: Organization: Hiroshima industrial promotion organization Participating institution and company: Kinki University (Project leader: Hideki Kyogoku) Imanishi Manufacturing Co., Ltd (Project sub-leader: Toshifumi Imanishi) Castem Co., Ltd (Leader: Hideki Nakayama) Hiroshima Prefectural Industrial Research Institute (Leader: Yasuhiro Sakamoto) Aspect Inc. (Leader: Masashi Hagiwara) The authors would like to thank the METI and MEXT of Japan for financial support ofthis research.The authors also would like to thank Prof. Bourell of the University of Texas at Austin for his invaluable suggestions, Toyo Aluminum K.K.

EEFORM FABRICATION OF BIOLOGICAL SCAFFOLDS BY PROJECTION PHOTOPOLYMERIZATIONLi-Hsin Han1, Gazell Mapili2, Shaochen Chen1, and Krishnendu Roy2Reviewed, accepted September 4, 2007ABSTRACTThis article presents a micro-manufacturing method for direct, projection printing of 3- dimensional (3D) scaffolds for applications in the field of tissue engineering by using a digital micro-mirror-array device (DMD) in a layer-by-layer process. Multi-layered scaffolds are microfabricated using curable materials through an ultraviolet (UV) photopolymerization process. The pre-patterned UV light is projected onto the photocurable polymer solution by creating the “photomask” design with graphic software. Poly (ethylene glycol) diacrylate (PEGDA), is mixed with a small amount of dye (0.3 wt %) to enhance the fabrication resolution of the scaffold.

Poly (ethylene glycol) diacrylate (PEGDA), is mixed with a small amount of dye (0.3 wt %) to enhance the fabrication resolution of the scaffold. The DMD fabrication system is equipped with a purging mechanism to prevent the accumulation of oligomer, which could interfere with the feature resolution of previously polymerized layers. The surfaces of the pre-designed, multi-layered scaffold are covalently conjugated with fibronectin for efficient cellular attachment. Our results show that murine marrow-derived progenitor cells successfully attached to fibronectin-modified scaffolds.INTRODUCTIONFreeform fabrication technology [1-10] has become a very useful tool for the manufacturing of biological scaffolds for tissue regeneration and stem cell engineering [10- 13] due to the efficient and simple process for creating three-dimensional (3D) microstructures.

Pre-designed, complex, 3D geometry in the micro- and nanometer scale can be created using polymeric materials without the use of lithographical masks or mechanical molds. The resolution given by DMD projection printing (DMD-PP) can create micro-pores comparable to the size of biological cells, i.e., micrometers. Depending on the number of pixels in a micromirror array, a DMD based system is capable of simultaneously forming structures in an area of several centimeters. The high resolution and fast fabrication speed make the DMD based system one of the most promising technology for fabricating tissue engineering scaffolds.In a typical process of DMD-based fabrication, a biological scaffold is formed on a moving-stage placed in a vat of photocurable monomer. The micro-mirror array from the DMD chip forms a pattern, created by graphic software, forming the geometry of a single layer for the 3D scaffold.

The micro-mirror array from the DMD chip forms a pattern, created by graphic software, forming the geometry of a single layer for the 3D scaffold. An ultraviolet (UV) light source illuminates the micromirror- array, and the illuminated pattern is projected, through an optical lens, onto the surface of the photo-curable monomer to form a single-layered structure. After one layer of monomer is selectively cured, the stage moves downward, as a new layer of monomer is formedabove the previously cured layer. The cross-section pattern can be changed in between scaffold layers, and the process is repeated until the entire scaffold is formed.The DMD-based fabrication process is, however, challenged by deep UV penetration in the solution and unwanted light-scattering. The curing depth of the UV light could be too deep such that previous layers are affected by the current photo-pattern being displayed.

The curing depth of the UV light could be too deep such that previous layers are affected by the current photo-pattern being displayed. UV-light scattering by previously formed microstructures may polymerize the monomer within the previous layers, causing a lost in feature resolution during the fabrication process. Lowering the UV intensity could fix these problems, but it could also degrade the resolution of the exposure layer. These problems became limiting factors to fabrication in our previous study [1].Scattering and penetration problems can be resolved by mixing UV dyes into photocurable monomers prior to irradiation. The function of the UV dye can be understood through Beer’s Law [14]. Assuming that a monomer has an absorption constantαand a UV light intensity,, the UV intensity becomes at a depth z to give the following:I0II=I0)αexp ( z ⋅−,(1)is a threshold intensity at which the monomer starts to cure.

CD is the curing depth, α becomes larger (α′ ) and CD becomes smaller CDI ( (CD’); the addition of the UV-dye to the monomer increases the absorption constantαand decreases the curing depth, and the structural resolution is therefore enhanced. Figures 1B and 1C show the patterned-polymerization of honeycomb-shaped structures with and without dye addition. In the absence of UV dye (Figure 1B), the monomer within the pores was partially cured by undesired scattering, and the structure had decreased resolution. With a small amount of dye in the same monomer solution, (Figure 1C), the structure is polymerized with increased resolution.Figure 1A shows the exponential decay of light intensity as z increases.I. When the UV dye is added,=)ICureAnother challenge encountered with the DMD system is the accumulation of oligomers within previously polymerized layers of the multi-layered construct.

When the UV dye is added,=)ICureAnother challenge encountered with the DMD system is the accumulation of oligomers within previously polymerized layers of the multi-layered construct. These oligomers are partially cured monomers, weakly cross-linked by unwanted scattered and penetrating UV light. To overcome this problem, it is necessary to purge the structure with new monomer solution between scaffold layers during the fabrication process.In this research, we built multilayered 3D structures for cellular behavior studies and for tissue engineering applications using the DMD-based fabrication system. The monomer that forms the scaffold was loaded with a small amount of UV absorbing dye to enhance the structural resolution. Our fabrication equipment also includes a simple purging system to prevent the accumulation of oligomers.

Our fabrication equipment also includes a simple purging system to prevent the accumulation of oligomers. Upon fabrication, scaffolds were biologicallyCureLight Intensity (A.U.)’D C.CAα′=.) .U A.( h t p e DD C.CAα=.(B)(A)100 µm(C)ICureI0Figure 1. The curing of monomer with and without the addition of the UV dye, under the same intensity of exposure, (A) Theoretical model: The curing depth CD decreases when the absorption constant is increased after the addition of dye. (B) Photo-patterning of poly(ethylene glycol) diacrylate on a glass slide. The pores of a honeycomb structure are sealed by unwanted curing caused by scattered light. (C) Patterning the same monomer with the addition of UV dye TINUVIN 234 at a concentration of 0.2wt%. The geometry of the honeycomb structure had increased feature resolution.modified for cell attachment.

The geometry of the honeycomb structure had increased feature resolution.modified for cell attachment. We show the intricate architecture of the 3D scaffold and also demonstrate efficient cell attachment within the porous tissue engineering construct.THE DMD PROJECTION PRINTING SYSTEMThis scaffold fabrication system (Figure 2A) consists of a vat containing the polymer, a servo-stage (CMA-25-CCCL & ESP300, Newport) that supports the construct, a syringe pump, a DMD system (Discovery 1100, Texas Instruments), a UV lamp (200W, S2000, EXFO), and a fixed-focal lens (Edmond Optics). The syringe pump is created by connecting a 3mL disposable syringe barrel to the servo-stage and is used to inject fresh photocurable monomer solution to an outlet located on the stage. The pump is automatically operated and is used in between layer fabrication. The DMD chip is composed of an array of micro-mirrors that project the patterned photomask onto the photocurable monomer.

The pump is automatically operated and is used in between layer fabrication. The DMD chip is composed of an array of micro-mirrors that project the patterned photomask onto the photocurable monomer. These micro-mirrors are illuminated by the UV light from the lamp using an 8mm light guide, and the fixed focal lens focuses and projects the images onto the translational stage. The moving-stage, syringe-pump, DMD system, and the UV lamp are connected to a personal computer, and the DMD-base system is controlled through a driver interface supplied by Texas Instruments.SCAFFOLD FABRICATIONPreparation of the Photocurable MonomerDMD ArrayUV lightFocus LensPEGDA/ MAAOutleServo- StageC6F14Chemical Bath(A)(B)(C)Cross-section Images:(1)(2)(3)Figure 2. The schematics of the DMD fabrication system showing the fabrication scheme (A-C), and the patterns of the scaffold cross-section (inset).

The schematics of the DMD fabrication system showing the fabrication scheme (A-C), and the patterns of the scaffold cross-section (inset). (A) The forming of one structural layer by a patterned UV- cross-section image. (B) A purging process to remove partially cured monomer. (C) Forming the next structural layer; the stage is positioned 50 um below the liquid surface to define the thickness of a new layer.Perflorohexane (99.5%) and poly (ethylene glycol) diacrylate (PEGDA, Mn 258) were purchased from Sigma-Aldrich and used as received. Methacrylic acid (MAA) from Sigma- Aldrich was distilled before use to remove inhibitors. The photo-initiator, Irgacure 2959, and UV dye, TINUVIN 234, were provided by Ciba Chemistry; both chemicals were used without further purification.To prepare the photocurable monomer solution, 10 wt% of Irgacure 2959 and 0.2 wt% of TINUVIN 234 were added into a 4:1 (volume ratio) mixture of PEGDA and MAA.

The monomer solution was sonicated for 30 minutes and degassed for 15 minutes.Pre-Designed Scaffold FabricationNew Layer(A)(B)(C)(D)Figure 3. SEM pictures (Figure 3A-C) and fluorescence (Figure 3D) micrographs show that the D1 cells attach and secrete extracellular matrix onto the surfaces of the fibronectin-modified scaffold.The polymer vat was filled with perfluorohexane (C F6 14), an inert chemical with high density (1.685 g/cm3) and a molecular polarity significantly lower than the photo-curable monomer. Because of these properties, the photocurable monomer forms a layer on top of perfluorohexane within the polymer vat. The majority of the stage is immersed in perfluorohexane. A key advantage of using an inert, heavy liquid like perfluorohexane is that only a small amount of photocurable monomer solution is necessary during fabrication as the perfluorohexane acts as a “filler” material for the bottom region of the stage.

The syringe pump was also filled with perfluorohexane to be used as the purging agent. The focal plane of the DMD-PP was determined, and the stage was set accordingly. The height of the polymer vat was adjusted to allow a thin layer of the photocurable monomer to form on the outlet of the syringe pump. A UV photomask from the DMD system selectively solidifies the monomer into a thin layer, creating the first layer of the structure. The shape of the outlet was properly designed, such that the scaffold layers completely cover the outlet.The power of the UV image was determined to be approximately 5mW/cm2, and UV exposure time was set for 60 seconds. After the first layer was built (Figure 2B), the stage moves the structure below the perfluorohexane/monomer interface, and the syringe pumps perfluorohexane through the microstructure on the outlet. The perfluorohexane solution pushes the uncured material (which includes oligomers) back to the monomer layer.

The perfluorohexane solution pushes the uncured material (which includes oligomers) back to the monomer layer. Afterthe purging step, the stage moves up and locates the top of the structure slightly higher than the perfluorohexane/monomer interface. The syringe aspirates monomer at a volume equal to the amount purged by perfluorohexane, and the structure is filled with fresh monomer. Following this step, the stage moves to 50 µm below the position where the previous layer was created (Figure 2C), and the next layer is then created. The patterned-polymerization and purging cycles are repeated until the entire, pre-designed scaffold is built.As shown in the inset of figure 2, we used three different cross-sectional images (1-3) to pattern the scaffold with hexagonal-shaped porosity. The images were used in a sequence: 1-1-1-2-1-1-1-3-1-1-1-2-1-1-1-3-1.

The images were used in a sequence: 1-1-1-2-1-1-1-3-1-1-1-2-1-1-1-3-1. The layered structure of the scaffold includes four honeycomb layers, neighboring honeycomb layers are spaced by three rectangular rims, and the space between two honeycomb layers was measured to be 150 µm.By using UV dye TINUVIN 234, we successfully adjusted the curing depth of monomer to about 50 µm. There is little scattering interference between the new layer and previously formed layer.

There is little scattering interference between the new layer and previously formed layer. Under a brightfield microscope, we found that the geometry of each of the honeycomb layers was clearly defined, and the pores of the honeycomb structures were free from unwanted polymerization or oligomer accumulation.Surface Modification of the Preformed Scaffold with FibronectinAfter the scaffold was fabricated, tethered carboxyl groups from the methacrylic acid (pre-mixed in the monomer) was activated using 1-ethyl- 3- [3-dimethylaminopropyl] carbodiimide hydrochloride/ N- hydroxysulfosuccinimide chemistry (EDC/ sulfo-NHS, in EDC Pierce Technologies, Rockford, IL). bioconjugation chemistries to create amide bonds. Carboxyl groups are converted to amine-reactive sulfo-NHS esters by EDC with the presence of sulfo-NHS.

bioconjugation chemistries to create amide bonds. Carboxyl groups are converted to amine-reactive sulfo-NHS esters by EDC with the presence of sulfo-NHS. 0.4M EDC and 0.4M sulfo-NHS in 0.1M MES buffer [(2-(N-morpholino) ethane sulfonic acid), pH 6.5], were added to the scaffolds at a total volume of 1.5 mL and incubated on a rotator for 2 h at room temperature. Using a low-affinity microcentrifuge tube, a 1.5 mL solution of fibronectin (10 g/mL) was then added to the scaffolds upon carboxyl activation, and incubated at room temperature for a 24 h period. Scaffolds were then sterilized using 70% ethanol for 30 minutes prior to cell seeding, and rinsed several times with PBS to remove unconjugated fibronectin and ethanol.is a commonly used crosslinkerMurine Marrow-derived Progenitor CellsA bone marrow progenitor cell line derived from mice, D1 ORL UVA (ATCC, Manassas, VA), was seeded onto the fibronectin-modified scaffolds created by the DMD- based system.

Prior to seeding the cells onto the scaffolds, an amine-reactive fluorescent dye, CellTrace™ Far Red DDAO-SE (Molecular Probes, Eugene, OR) was used to stain the cells red, following the manufacturer’s protocol. Scaffolds were placed on sterilized parafilm within a tissue culture plate and then suspended in primary medium that contained 250,000 D1 cells. The suspended cell solution forms a “ball” on top of the scaffold, therebylocalizing the cells onto the fibronectin-modified scaffold, due to the hydrophobic surface of the parafilm. After a 4 h seeding period, the rest of the primary medium was added to the tissue culture plate. Primary medium used to culture D1 was composed of 10% w/v FBS (ATCC) and 1% w/v penicillin streptomyocin in Dulbecco’s Modified Eagle’s Medium (ATCC). The cell-scaffold construct was placed in a humidified incubator (5% CO2, 37°C) and kept in culture for a 48 h period before fixing in a 10% formalin solution.

The cell-scaffold construct was placed in a humidified incubator (5% CO2, 37°C) and kept in culture for a 48 h period before fixing in a 10% formalin solution. Scaffolds were then prepared for scanning electron and fluorescence confocal microscopic analyses.Scanning electron (Figure 3A-C) and fluorescence (Figure 3D) micrographs show that D1 cells attach and secrete extracellular matrix onto the surfaces of the fibronectin- modified scaffold. Fibronectin is a highly characterized protein known to bind to cell integrins, causing cell anchorage onto surfaces bound to this particular protein. Figure 3D is a three-dimensional compiled fluorescence micrograph of attached cells created from individual images obtained through confocal microscopy.

Figure 3D is a three-dimensional compiled fluorescence micrograph of attached cells created from individual images obtained through confocal microscopy. The scaffold depicted here is composed of four layers with a wall thickness of 50 µm and a hexagonal pore geometry of approximately 150 µm.CONCLUSIONA porous, multilayered, 3D scaffold for studies in cellular behavior was successfully created by patterning a photocurable monomer using a DMD-based fabrication tool. The monomer includes a small amount of absorbing dye (TINUVIN 234, 0.2wt%) which enhances the geometric resolution of the scaffold and reduces the curing depth of the monomer. The fabrication tool also includes a perfluorohexane-based purging mechanism to remove oligomers within the pores of the curing structure. The surfaces of the 3D scaffolds were biochemically modified with fibronectin for efficient cellular attachment.

The surfaces of the 3D scaffolds were biochemically modified with fibronectin for efficient cellular attachment. Data from SEM and confocal fluorescence micrographs confirm the successful seeding, attachment, and proliferation of D1 cells onto these pre-designed, micro-fabricated porous structures.ACKNOWLEDGMENTSThis work is supported by a Young Investigator Award from the Office of Naval Research and a grant (CMMI 0600104) from the National Science Foundation to S.C. We appreciate the donation of the DMD tool kit from Texas Instruments. We also thank the computer support from Intel’s Higher Education Program.REFERENCES[1] Lu, Y., Mapili, G., Suhali, G., Chen, S.C., and Roy, K. 2006, "A Digital Micro-mirror Device (DMD)-based System for the microfabrication of Complex, Spatially Patterned Tissue Engineering Scaffolds," J. Biomed. Mater. Res A, 77A (2), pp. 396-405.[2] Sun, C., Fang, N., Wu, D.

J., Van Stryland, E.W., Chapela, V. M., and Percino, J., 2000, “Synthesis and characterization of novel rigid two- photon absorbing polymers,” J. Am. Chem. Soc. 122, pp. 1217.[6] Nguyen, L. H., Straub, M., and Gu, M., 2005, "Acrylate-Based Photopolymer for Two- Photon Microfabrication and Photonic Applications," Adv. Funct. Mater., 15 (2), No.2, pp. 209.[7] Straub, M., Nguyen, L. H., Fazlic, A., and Gu, M., 2004, "Complex-shaped three- dimensional microstructures and photonic crystals generated in a polysiloxane polymer by two-photon microstereolithography," Opt. Mater., 27, pp. 359–364.[8] Vozzi, G., Flaim, C., Ahluwalia, A., and Bhatia, S., 2003, "Fabrication of PLGA scaffolds using soft lithography and microsyringe deposition," Biomaterials, 24, pp. 2533– 2540.[9] Itoga. K., Yamato, M., Kobayashi, J., Kikuchi, A., and Okano, T., 2004, "Cell micropatterning using photopolymerization with a liquid crystal device commercial projector," Biomaterials, 25, pp.

K., Yamato, M., Kobayashi, J., Kikuchi, A., and Okano, T., 2004, "Cell micropatterning using photopolymerization with a liquid crystal device commercial projector," Biomaterials, 25, pp. 2047–2053.[10] Sachlos, E. and Czernuszka, J. T., 2003, “Making tissue engineering scaffolds work. Review: the application of solid freeform fabricationtechnology to the production of tissue engineering scaffolds,” Eur. Cell. Mater., 5, pp. 29 –39; discussion pp. 39 – 40.[11] Griffith, L. G., 2002, “Emerging design principles in biomaterials and scaffolds for tissue engineering,” Ann. N. Y. Acad. Sci., 961, pp. 83-95.[12] Orban, J. M., Marra, K. G., and Hollinger, J. O., 2002, “Composition options for tissue- engineered bone,” Tissue Eng. 8, pp. 529 –539.[13] Sharma, B. and Elisseeff, J. H., 2004, “Engineering structurally organized cartilage and bone tissues,” Ann. Biomed. Eng. 32, pp. 148-159.[14] Brown, B., Foote, C.

eeform Fabrication of Electroactive Polymer Actuators and Electromechanical DevicesEvan Malone, Hod LipsonMechanical and Aerospace Engineering, Cornell University evan.malone@cornell.edu Reviewed, accepted September 1, 2004Abstract In pursuit of the goal of producing complete electromechanical systems entirely via solid freeform fabrication, we are developing a library of mutually compatible, functional, freeform elements. Several essential elements – actuation, sensing, and control electronics - still remain to be incorporated into this library. Conducting polymers (CP) are a class of materials which can be used to produce all of these functionalities. Meanwhile, research into actuatable “smart” materials has produced other candidate materials for freeform fabricated actuators that are compatible with our library.

Meanwhile, research into actuatable “smart” materials has produced other candidate materials for freeform fabricated actuators that are compatible with our library. We have succeeded in manually producing air-operable actuators that have processing and operating requirements that are compatible with our power source and mechanical component library elements. A survey of candidate actuator materials is presented, experiments performed with two types of actuator materials are described, and complete SFF-producible actuator devices are demonstrated.IntroductionImagine a compact fabrication system which could print out a complete working device, including power sources, electronics, actuators, sensors, and mechanical components, even biological materials, in almost any geometry and interconnection. The devices produced would not appear the same or operate on the same principles as devices built from mass produced components.

The devices produced would not appear the same or operate on the same principles as devices built from mass produced components. Whole new classes of unforeseen functional devices and products would be likely to emerge from the ability of such a tool to decouple geometry and functionality to a greater extent than ever before. There are different sets of mutually compatible materials and processes which might be used to create such a library and a system. We have chosen to focus primarily on polymers, and low temperature (<300C) processes. We have also limited our system to liquid state processes, either through fused deposition, or robotic dispensing from syringe. Within these restrictions, we have been working steadily to add new freeform functionalities while continuing to improve the quality, performance, and geometry independence of the batteries, structures, joints and wiring already within the library (Malone, Rasa et al. 2004).

2004). It is the goal of the current work to add the functionality of actuation to this library. Actuator materials have been identified which are amenable to our freeform fabrication processes, and actuator designs have been successfully demonstrated which can be freeform fabricated with and can operate in concert with the other functionalities.Actuator materialsIn the interest of producing biomimetic robotics and smart structures, and for medical applications, there is significant research into electromechanically active materials – materials which change shape, volume, modulus, or some other mechanical property in response to some kind of controllable signal. The number of such actuator materials is increasing quickly and the field of research is building momentum.

The number of such actuator materials is increasing quickly and the field of research is building momentum. The majority of materials investigated still require very specialized processing which prevents them from being cofabricated with other functionalities.697Nevertheless, several types of actuatable materials exist which may be compatible with our processing regime (Table 1).Table 1Material TypeCPPi-stackingIPMCDielectric elastomer Ferroelectric (piezoelectric) polymer Liquid-crystal elastomer Hydrogel Nanotube matActuation Signal1V10V2V>50V/um>10V/umthermal (80C- >100C) 3V 1.2Actuation Strain (%)520bending>100240>10 0.04Actuation Stress (MPa)5<321 (theoretical)0.1-7160.2>10Actuation Bandwidth (Hz) 11020k>1k10<0.1 >1kElectro- mechanical Efficiency (%) 0.1< (theoretical)66>80 (theoretical)References(Madden, Madden et al. 2001; Smela 2003) (Anquetil, Yu et al.

2001; Smela 2003) (Anquetil, Yu et al. 2002) (Shahinpoor and Kim 2001) (Pelrine, Kornbluh et al. 2000) (Zhang, Li et al. 2002)(Thomsen, Keller et al. 2001) (Liu and Calvert 2000) (Baughman, Cui et al. 1999; Lewis, Kim et al. 2000; Spinks, Wallace et al. 2001)Given that it is our goal to produce a library of mutually compatible functionalities, there are some further constraints on the choice of materials and actuator implementations.1. We desire that the actuator provide useful output using only power provided by Zn-air batteries which we can produce. The baseline performance of our 5g (active materials) cells is about 10mW at 1V continuous output for >50 hours. Although we currently lack the ability to fabricate electronics to perform voltage level conversion, we can conceive of placing a few cells in series and/or in parallel to meet power requirements.

Although we currently lack the ability to fabricate electronics to perform voltage level conversion, we can conceive of placing a few cells in series and/or in parallel to meet power requirements. Dissipation within the cells will likely limit series multi-cell battery output to < 10V – mass may limit the number of cells in parallel, depending on the desired end product (e.g. a mobile robot). Hence we require that useful actuation can be achieved at 10V or below.2. The mechanical components that we can currently produce are of relatively low stiffness, so for efficiency of mechanical coupling, we prefer actuator materials which have large strain output to those that high stress output, but low strain.3.

We require that the actuator operate in air, under ambient conditions.Referring to Table 1 above, these requirements allow us to eliminate several of the candidate material types.Pi-stacking material films are currently only produced electrochemically, so no obvious route to freeform fabrication yet exists.Dielectric elastomers are very attractive electromechanically, but the high actuation field would require producing and electrically and mechanically coupling to elastomer films as thin as 200nm, which seems extremely daunting given our fabrication processes.Piezoelectric polymers require high fields (very thin films), and typically require some solid-state processing (poling) to become electromechanically active.698Liquid crystal elastomers are thermally actuated, and although this could be achieved through resistive heating, the additional complication and likely low efficiency make them less appealing.Hydrogels have been produced via ink-jet (Yoshioka and Calvert 2001), but are delicate materials, slow to actuate, and suffer limited active life because of dehydration and leakage during actuation.Nanotube mats have very high stress output and efficiency, operate at low voltage, but have low strain output.We have explored the remaining two material types, CP and IPMC in greater detail.Conducting Polymer Actuators Conducting polymers are long chain organic molecules in which successive carbon atoms along the axis of the chain are bound alternately by one shared electron pair (single, or σ-bond) and by two shared electron pairs (double, or π-bond).

Like most organic polymers, CP’s are normally poor electrical conductors. By adding or removing electrons from the CP chains, and embedding positive or negative ions, respectively, into the chains to maintain electrical neutrality, it is possible to make the locations of the π-bonds less well defined (MacDiarmid 2001). Classically speaking, this allows some of the electrons more freedom to move from atom to atom, increasing electrical conductivity. Through this process known as redox doping, the conductivity of CP’s can be varied continuously from insulating to semi-conducting to metallic. This allowse-+V--V+V--VOxidatione-ReductionPolymer doped with large anionsFree ions in electrolyteFree ions in electrolyteCations move into polymer chains to compensate for gained electronsFigure 1. Illustration of the conformational change in a CP caused by ion diffusion.

Illustration of the conformational change in a CP caused by ion diffusion. Conformational change results in volume change and mechanical actuation.the fabrication in CP of many of the electronic devices currently made with doped silicon. Conducting polymers change their volume when redox doping occurs. If the CP is placed in contact with an electrolyte (salt solution with free ions) and an electrical potential is applied between the polymer and another electrode immersed in the electrolyte, free ions (and attached solvent molecules) from the electrolyte diffuse into or out from between the tangled polymer chains to compensate for the changed charge of the polymer, thereby causing the overall volume of the polymer to change (Figure 1) (Bar-Cohen 2001). The amount, rate, and polarity of the volume change depend on specific sizes and charges of the dopant ions and the electrolyte solvent used.

2001). Small strain responses are observable at frequencies > 30Hz. There is some indication that out of plane actuation strains of 30% or more are achievable (Smela and Gadegaard 2001), and there is some preliminary work on novel actuator device geometries which might exploit this large strain (Della Santa, De Rossi et al. 1997). Unfortunately, practical electromechanical efficiency is currently < 1%, although Madden (Madden, Madden et al. 2001) suggests some routes to improvement. The sensitivity of CP actuator performance to the choice of electrolyte, because of ion size and mobility and electrolyte stability, means that electrolytes and CP materials must be explored in combination. The majority of CP research has involved liquid electrolytes, because liquids allow high ion mobility, but a few air-operable CP actuator designs have been investigated (Vidal, Plesse et al.; Madden, Cush et al. 1999; Hutchison, Lewis et al.

1999; Hutchison, Lewis et al. 2000), and provide a good starting point for the selection or development of non-liquid electrolytes or encapsulation methods.PotentiostatVElectrodeCounterElectrodeWorkingGrowing PolymerFilmElectrodeReferenceMonomer, Solvent, Dopant Electrochemical Polymerization(b) Figure 2. (a) Electrochemical polymerization apparatus, (b) successful actuation of electrochemically produced polypyrrole film, and (c) poor mechanical properties of dispersion-cast polypyrrole(a)(c)Conducting polymers are very appealing as actuator materials from the standpoint that electronics and sensors made with conducting polymers are already being produced using printing techniques, and are rapidly being commercialized, suggesting that facility with freeform fabrication of CP materials might accelerate the completion of our library of functionalities.

The vast majority of work on CP actuation uses materials and processes that are not readily adaptable to SFF. The most common CP actuator material is polypyrrole, and the best published actuation results (Madden, Madden et al. 2001) involve polypyrrole produced via low-temperature (-30C) electrochemical synthesis (Figure 2a), which involves using electric fields in a solution of pyrrole monomer and electrolyte to cause oxidative polymerization.Very high molecular weight films of polypyrrole, already doped and with good mechanical properties can be produced this way. As a baseline experiment, and to build experience with these materials, we produced some polypyrrole electrochemically following the method of (Yamaura, Hagiwara et al. 1988), and demonstrated its actuation in liquid electrolyte (Figure 2b).Despite the low efficiency mentioned in Table 1, air operable CP actuator designs have been published with performance sufficient to warrant some exploration.

The key challenge is to700find an SFF-amenable means of producing conducting polymer films with high molecular weight (good mechanical properties), and high conductivity. As a rule of thumb, conducting polymers are not melt processable or soluble in their conducting (doped) state, although they may be so in the insulating state. Polypyrrole and other conducting polymers are available in the doped form as dispersions (small particles suspended in liquid medium) in water or other solvents. We investigated the mechanical properties of films produced from these materials by casting films of polypyrrole dispersion (Sigma-Aldrich), and of PEDOT/PSS (Sigma-Aldrich). The results were unsatisfactory with polypyrrole (Figure 2c), while PEDOT/PSS can produce fairly robust and flexible films.

The results were unsatisfactory with polypyrrole (Figure 2c), while PEDOT/PSS can produce fairly robust and flexible films. To achieve air-operable actuators, some kind of electrolyte must be used – either a “solid polymer electrolyte” (SPE) which contains some liquid to allow ion migration, or a liquid electrolyte surrounded by some kind of encapsulation. Unfortunately, films cast from dispersions of these two materials remain very sensitive to the liquid components of the electrolyte, and no successful CP/electrolyte combinations were found.1cm1cmFigure 3. Actuation of solution-cast P3OT in liquid electrolyte; 5V step, elapsed time 5sIn (Dahman 1999) a method is described which allows a compromise on the solubility/conductivity tradeoff with two conducting polymers, P3OT (poly(3-octylthiophene- 2,5-diyl) and PANI (polyaniline).

In the processing method described, undoped P3OT or PANI powder is dissolved/dispersed in a solvent, then doped in solution with two separate dopants, one which improves the conductivity of the material, and one which improves the processability of the material. We produced a sample of P3OT, doped only with DBSA (dodecylbenzenesulfonic acid) as processing dopant, in toluene to explore the mechanical properties of the film. Films were cast from this solution onto a Teflon substrate, and allowed to air dry. When peeled off of the Teflon, the films were found to be flexible and elastic. The actuation of this material in a liquid electrolyte was verified clearly – the motion of the film is fast enough and large enough to observe with ease in video. Figure 3 shows the change in curvature with a 5V step over 5 seconds.Subsequently, an attempt was made to test actuation in air. A solid polymer electrolyte was produced using a modification of the methods in (Sansiñena, Olazábal et al.

A solid polymer electrolyte was produced using a modification of the methods in (Sansiñena, Olazábal et al. 1997). A low vapor pressure solvent, propylene carbonate was used to dissolve Hydrin (poly(epichlorohydrin- co-ethylene oxide, Zeon Chemical) rubber to act as a matrix for in order slow down the drying out of the electrolyte. A three-layer sandwich (P3OT-Hydrin SPE-P3OT) was made by spreading SPE on the surfaces of two 10mm X 40mm strips of P3OT film, then laying these face to face, SPE facing inward (Figure 4). Electrical contact was made via aluminum foil strips mounted on adhesive tape, which was then attached to the P3OT strips. A 1.5V DC signal at about 10mA was supplied. A very slow movement was observed to take place over about 4 minutes (Figure7015). The very slow actuation may be in part due to the low conductivity of the P3OT, which was prepared without the conductivity enhancing dopant recommended by Dahman.

The very slow actuation may be in part due to the low conductivity of the P3OT, which was prepared without the conductivity enhancing dopant recommended by Dahman. It is possible that the material is not actuating by ion diffusion, but instead by solvent electrolysis. A subsequent experiment will explore these possibilities.Conducting PolymerGel- or Solid Polymer Electrolyte3-Layer, Bending-beam ActuatorFigure 4. Schematic of three-layer, air operable CP actuatorFigure 5. Actuation of P3OT with SPE in air (curling of tip); 1.5V step, 4 minutes totalNo attempt was made to optimize the formulation of the SPE for this preliminary experiment. Significantly improved actuation can be expected from improvement in the P3OT doping and SPE formulation. Although (Chen and Inganas 1995) reports actuation of P3OT in liquid electrolyte, the present work is the first reporting of an air-operable CP actuator produced entirely by casting films from solutions.

The materials are all syringe-dispensible, and hence suitable for use in our freeform fabrication system.Ionomeric Polymer-metal Composite ActuatorsIonomeric Polymer-Metal Composites (IPMC’s) are actuators constructed from an anionic perfluorinated polymer membrane (e.g. a modified PTFE (Teflon) with perfluorinated sulfonate anion side groups, Figure 6) with surface electrodes. These materials are the key component in proton exchange membrane (PEM) fuel cells, and are commercially available (e.g. as DuPont Nafion). The materials are typically produced in what is called “acid form,” meaning that a proton is ionically bound to the sulfonate anions in the polymer side chains. The standard approach to IPMC fabrication (Kim and Shahinpoor 2003) involves purchasing a solid membrane, and replacing the proton by another cation (e.g. Li+) to improve actuation properties (Figure 6).

Li+) to improve actuation properties (Figure 6). The surfaces of the membrane are given metallic electrodes, typically by soaking in a metal salt solution, then chemical reduction of the metal salt.A voltage applied across the electroded surfaces of the composite causes the material to bend (Figure 7). This probably results from the migration of cations and solvent (the anionic side groups on the polymer being immobile) within the material, as well as from several associated effects: a stiffness change in the polymer, a change in electrostatic forces between the anionic side groups in the polymer, reorientation of the solvent molecules (Shahinpoor and Kim 2001). If the actuation signal is rapidly applied (a step), there is a rapid bending toward the negative electrode, followed by a slower relaxation back toward the positive electrode, resulting in a net motion toward the positive electrode.

The fast processes are probably cation motion, stiffness, polymer conformation, and solvent orientation related, while the slow process is probably solvent diffusing to equalize pressure through the material (Nernat-Nasser and Li 2002). A gradually ramped stimulus may result in only the slow processes occurring (Bar-Cohen 2001). IPMC actuators are superior to CP actuators in terms of their actuation speed and probably their702efficiency (difficult to find reported values in literature). Unfortunately, because their motion is not the result of net volume change, but only of bending, they are somewhat difficult to couple mechanically.+++-+++ - ++ -++++ ++Mobile cationSolvent moleculeAnionic side-chainFigure 6. Structure of DuPont Nafion polymer; exchange of cation (from H+ to Li+) to improve actuation; Rf is proprietary fluorinated groupFigure 7.

Structure of DuPont Nafion polymer; exchange of cation (from H+ to Li+) to improve actuation; Rf is proprietary fluorinated groupFigure 7. Possible IPMC actuation mechanism: pressure differential from electric field driven ion and solvent motionProduction of IPMC via SFF will enable the exploration of the effect of complex device geometry on actuator properties, as well as allowing the embedding of actuators within other devices, but the standard method of producing IPMC described above is incompatible with freeform fabrication. (Kim and Shahinpoor 2002) describes a method for producing IPMC actuators from “liquid Nafion,” which is a dispersion of Nafion particles in solvent, but note that the authors have not demonstrated freeform fabricated actuators. The described process has six steps, but, as noted below, one step is incompatible with freeform fabrication, and several others are problematic from the perspective of co-fabrication with other functionalities.1.

Obtaining liquid Nafion: Nafion dispersions are commercially available in several loading (5% - 20%) and medium combinations (generally water and/or alcohols). The dispersions are typically used to repair Nafion membranes which are being used e.g. in fuel cells. (Kim and Shahinpoor 2002) also refers to methods for production of Nafion dispersion from commercially available Nafion membranes, and a means of producing melt-processable Nafion. Converting from Nafion from H+ to Li+ cation: Typically, this step is carried out by 2. soaking a solid Nafion membrane in aqueous salt solution (e.g. LiCl) at elevated temperature for several days. This step is incompatible with freeform fabrication since it is performed on a solid membrane.3. Casting of film: (Kim and Shahinpoor 2002) observes that films cast directly from purchased Nafion dispersions have poor mechanical properties.

Casting of film: (Kim and Shahinpoor 2002) observes that films cast directly from purchased Nafion dispersions have poor mechanical properties. Improved films are achieved by adding DMF (dimethylformamide), a low vapor pressure solvent, to their dispersions. The amount of DMF added is not reported. Thicker films can be achieved through repeated drying and readdition of liquid. The authors cast material into a Petri dish. This step is problematic for freeform fabrication in that the dispersions need to be confined during drying to control film geometry and to prevent contamination of other fabricated components by the somewhat hazardous solvents involved.4. Annealing of film: The air dried films thus produced are still easily resuspended in solvent. We have verified this – within moments of immersing the film in deionized water, it disintegrates.

Annealing of film: The air dried films thus produced are still easily resuspended in solvent. We have verified this – within moments of immersing the film in deionized water, it disintegrates. This is undesirable because the film must be hydrated in order to actuate.703Annealing of the film is suggested in (Kim and Shahinpoor 2002), as well as by DuPont technical support, as a means of causing some degree of crystalization and a reduction of solvent susceptibility. The latter source recommends 15 minutes at 135C, while the former fail to suggest a duration, but mention 70C followed by 150C. The annealing process is incompatible with the goal of cofabricating actuators with other fucntionalities; batteries in particular would suffer some degradation as a result of this process.5. Electroding of film: The surfaces of the film are typically made electrically conducting by the metal salt reduction technique mentioned above, which is not SFF amenable.

Electroding of film: The surfaces of the film are typically made electrically conducting by the metal salt reduction technique mentioned above, which is not SFF amenable. Although not clearly demonstrated, (Kim and Shahinpoor 2002) does suggest a method of electroding that is. A metal (Ag, Pt, etc.) powder is suspended in a quantity of Nafion dispersion, then this material is cast as the first (bottom) layer, and the last (top) layer of a Nafion film.6. Hydration of film: The electroded film is soaked in deionized water for a few minutes to saturate it. Water leaks and evaporates from the devices during use, limiting operational life; some means of encapsulation is desirable to prolong in-air operation.We performed a series of experiments in order to modify or eliminate the problematic steps (steps 2, 3, 4, 6) in order to improve the compatibility of IPMC fabrication processes with our processing regime and requirements of other library functionalities.

For these experiments we purchased 10 wt. % aqueous Nafion dispersion from Sigma Aldrich. The dispersion is in H+ (acid) form, as purchased. Additional experiments were performed to verify published experiments and to explore other avenues.Conversion of cation (modification of step 2)To eliminate the problems with step 2, we achieve the exchange to Li+ form entirely in the liquid state through dialysis in two different processes. Consultation with DuPont Nafion technical support revealed that the aggregated particles in the Nafion dispersions are typically of at least 100kg/mol. For this reason we selected 25kDa pore size dialysis tubing as a compromise between Nafion retention and dialysis rate. In the first method, the Nafion dispersion is placed in the dialysis tubing, and the tubing is immersed in 1.5M LiCl solution. The solution is stirred continuously, and refreshed every 12 hours for six cycles.

The solution is stirred continuously, and refreshed every 12 hours for six cycles. Excess LiCl is then removed by immersing the dialysis tubing in deionized water, and similarly stirring and refreshing for three additional 12 hour cycles. The second, faster approach involves dissolving an excess of LiCl directly in the as-purchased aqueous Nafion, then dialysing it against deionized water for three 12 hour cycles. It must be noted that the weight percent of Nafion in the dispersions following the conversion to Li+ form via dialysis are not necessarily the same as in the purchased material.Casting of films/annealing (elimination of step 4)We explored several alternatives for the preparation of films. In general films were cast onto a Teflon substrate, and allowed to dry at room temperature under a fume hood (constant air flow). Initially, films were cast directly from the Li+ aqueous dispersion, without additives.

Initially, films were cast directly from the Li+ aqueous dispersion, without additives. Our experience replicated (Kim and Shahinpoor 2002), in that the films cracked upon drying, and the pieces were too fragile to handle.As a verification of published approaches, a ratio of 1:4, DMF:Nafion (post-dialysis) dispersion by volume resulted in transparent, durable films which can be peeled off of the substrate with tweezers and handled. The film was annealed (15 minutes at 135C), electroded by coating the surfaces with silver grease (CircuitWorks CW7100), hydrated in deionized water for704several minutes, and tested for actuation. The test was successful, with the device moving visibly and repeatedly within a second in response to variation of applied voltage between 0V and 2V, and current typically less than 10mA.

The test was successful, with the device moving visibly and repeatedly within a second in response to variation of applied voltage between 0V and 2V, and current typically less than 10mA. Application of 10V caused a short circuit failure of the device before video documentation could be obtained.To explore the effect of the DMF/water ratio, a quantity of aqueous dispersion was allowed to completely dehydrate. The resultant dry material weighed 1.4g. It was dissolved in 6 mL of DMF, resulting in a viscous, but transparent dispersion at 25wt%. Film was successfully cast from this concentrated dispersion, and was durable enough to handle. The film was prepared for testing exactly as above. This test was a complete failure, however, with no motion observed, regardless of applied voltage.

The film was prepared for testing exactly as above. This test was a complete failure, however, with no motion observed, regardless of applied voltage. Little or no current flowed, so electrical shorting was not the cause.In all of the above cases, the material was still found to require annealing, and to be somewhat fragile, especially after hydration. To remedy this problem, a small quantity of Hydrin SPE matrix rubber dissolved in DMF (about 10 wt% Hydrin in DMF) was added to the Nafion dispersion. The idea is that the very long chain Hydrin polymer molecules might act as a reinforcing matrix for the Nafion without hindering ion and solvent transport. 0.1g of dehydrated Nafion was mixed with 2mL of DMF and this was mixed with 0.1g of Hydrin/DMF solution. The resultant material was cast, and the resultant films was found to be mechanically superior to Nafion alone.

The resultant material was cast, and the resultant films was found to be mechanically superior to Nafion alone. A piece of this material was hydrated without annealing, and found to maintain its integrity about as well as annealed Nafion. When tested, the material was not found to be electromechanically active.Figure 8. Actuation of annealed Nafion/Hydrin blend; 5V step, elapsed time 45sFigure 9. Actuation of unannealed Nafion/Hydrin blend; 5V step, elapsed time 60sThis failure suggested that either the Hydrin was interfering with the actuation of the Nafion, that annealing was essential to actuation, or that the lack of water during casting was at fault. In that this last issue was in common between the Hydrin/Nafion blend and the concentrated Nafion cast from straight DMF, both of which failed to actuate, the Hydrin/Nafion blend (1.5mL ) was tried with the addition of aqueous Nafion dispersion (4mL).

The liquid clouded immediately upon addition of the aqueous dispersion, and the cast film was opaque, but retained the good mechanical properties even upon hydration without annealing. A piece of this705film was annealed and tested for actuation as before, with silver grease electrodes. This test was successful (Figure 8), and a subsequent test also verified the actuation of unannealed material (Figure 9). With this polymer blend we have achieved an IPMC that does not require annealing, although annealing still seems to be beneficial for the mechanical properties of the material.Electroding (exploration, verification of published approach)A variety of materials were explored for electroding, including silver grease, conducting polymers (including the PEDOT/PSS and P3OT used in CP actuator experiments), and metal poweders.

A series of experiments were performed in which the electrode materials were cast or laid down on a Teflon substrate, and Nafion dispersion or Nafion/Hydrin blend was cast on top of them. The majority of these experiments failed, although the reasons seemed to be one of two: chemical incompatibility (with P3OT; likely the DBSA dopant or residual toluene was the cause), or contraction and crumpling of the electrode layer during evaporation of solvent from the subsequent Nafion or Nafion/Hydrin layers. There were two successes. A thick layer of PEDOT/PSS about 0.25mm, which was fully air dried was able to survive having Nafion cast ontop of it, but did crack somewhat. The method suggested by (Kim and Shahinpoor 2002) seems to work best. 0.5g of 2µm Ag powder was mixed with 250µL of aqueous Nafion dispersion, and ~10µL of DuPont Zonyl FSO-100 fluorosurfactant was added to promote dispersion of the silver.

0.5g of 2µm Ag powder was mixed with 250µL of aqueous Nafion dispersion, and ~10µL of DuPont Zonyl FSO-100 fluorosurfactant was added to promote dispersion of the silver. This material was cast as a film, and was found to form a high conductivity material via digital multimeter.Encapsulation and hydration (modification of steps 3 and 6)Standard IPMC electroding techniques result in porous electrodes. The result is that water leaks out of the actuator, as well as evaporating from it. Not only does this reduce the operating life of the device (at least without rehydration), but it has been suggested that leakage reduces the actuation efficiency of IPMC devices by reducing the achievable internal pressure gradients in the material. For this reason an impermeable film electrode is desirable, and is why conducting polymers were explored as electrode materials. Unfortunately, PEDOT/PSS, which was marginally successful, is also water soluble.

Unfortunately, PEDOT/PSS, which was marginally successful, is also water soluble. Applying an impermeable, low modulus skin to the IPMC was evaluated as a means of preventing leakage, regardless of the choice of electrode. The candidate membrane material selected is PDMS (poly(dimethylsiloxane), Dow Corning 734 Flowable Silicone RTV rubber). This material is within the suite of materials used in our SFF system; we routinely deposit 0.3mm thick films of this material. A sample of Nafion/Hydrin film was cast and air dried, and painted with PEDOT/PSS aqueous dispersion. Rather than hydrating by soaking the film, which would likely damage the PEDOT, it was assumed that the water from the PEDOT would serve to hydrate. When the PEDOT was visibly dry, electrical contact was made via wires with dots of silver grease, and the sample was mounted on Scotch tape. The entire actuator was then coated by hand in PDMS. Once dry, the device was tested for actuation.

The entire actuator was then coated by hand in PDMS. Once dry, the device was tested for actuation. Figure 10 shows the device tip elevating about 1mm in 50s in response to 5V.While the performance is not spectacular, actuation was definite and repeatable, validating the approach of hydration via PEDOT, and the compatibility of silicone encapsulant with actuation. Current was less than 10mA. From the perspective of producing these actuators via SFF, especially in unusual geometries and integrated with other functionalities, this successful test of silicone as an encapsulant is critical: silicone can be used to confine the liquid actuator materials during film casting, is electrically insulating, inert and chemically resistant,706but can be made thin enough, and has a sufficiently low modulus to not interfere excessively with actuation.Figure 10.

Actuation of silicone encapsulated Nafion/Hydrin; 5V step, elapsed time 50sConclusionIn the quest to develop a complete library of functional electromechanical elements which can be freeform fabricated and are mutually compatible, we have been investigating electromechanically active materials. Conducting polymers and proton exchange membrane materials have been identified as most compatible with our liquid deposition/fused deposition processes and the other functional elements which we can already produce. Significant new processing methods and material modifications have been developed. Actuators of both types have been made by hand-dispensing materials from syringes, layer by layer, which is a reasonable proxy for robotic freeform fabrication, and in-air actuation has been demonstrated conclusively in at least one case. Producing an isolated actuator with our SFF system using either of these materials will not pose a significant challenge.

Producing an isolated actuator with our SFF system using either of these materials will not pose a significant challenge. The major tasks ahead are the optimization of materials and processes, the electromechanical performance of these devices, and the development of designs and methods which will allow the co-fabrication of actuators with other functionalities in an integrated, operational device.the significant challenge of measuringAcknowledgments This work was supported in parts by the U.S. Department of Energy, grant DE-FG02- 01ER45902.References Anquetil, P. A., H.-h. Yu, et al. (2002). Thiophene Based Molecular Actuators. Smart Structures and Materials 2002:Electroactive Polymer Actuators and Devices (EAPAD), San Diego, CA, USA, SPIE Press.Bar-Cohen, Y. (2001). Electroactive Polymer (EAP) Actuators as Artificial Muscles - Reality, Potential andChallenges, SPIE Press.Baughman, R. H., C. Cui, et al. (1999). "Carbon Nanotube Actuators." Science 284(5418): 1340-1344.

eeform Fabrication of Ionomeric Polymer-Metal Composite ActuatorsEvan Malone, Hod LipsonMechanical and Aerospace Engineering, Cornell University evan.malone@cornell.eduReviewed, accepted August 3, 2005AbstractIonomeric polymer-metal composite (IPMC) actuators are a type of soft electromechanically active material which offers large displacement, rapid motion with only ~1V stimulus. IPMC’s are entering commercial applications in toys (Ashley 2003) and biomedical devices (Soltanpour 2001; Shahinpoor 2002; Shahinpoor, Shahinpoor et al. 2003; Soltanpour and Shahinpoor 2003; Soltanpour and Shahinpoor 2004), but unfortunately they can only actuate by bending, limiting their utility. Freeform fabrication offers a possible means of producing IPMC with novel geometry and/or tightly integrated with mechanisms which can yield linear or more complex motion.

Freeform fabrication offers a possible means of producing IPMC with novel geometry and/or tightly integrated with mechanisms which can yield linear or more complex motion. We have developed materials and processes which allow us to freeform fabricate complete IPMC actuators and their fabrication substrate which will allow integration within other freeform fabricated devices. We have produced simple IPMC’s using our multiple material freeform fabrication system, and have demonstrated operation in air for more than 40 minutes and 256 bidirectional actuation cycles. The output stress scaled to input power is two orders of magnitude inferior to that of the best reported performance for devices produced in the traditional manner, but only slightly inferior to devices produced in a more similar manner. Possible explanations and paths to improvement are presented.

Possible explanations and paths to improvement are presented. Freeform fabrication of complete electroactive polymer actuators in unusual geometries, with tailored actuation behavior, and integrated with other freeform fabricated active components, will enable advances in biomedical device engineering, biologically inspired robotics, and other fields. This work constitutes the first demonstration of complete, functional, IPMC actuators produced entirely by freeform fabrication.IntroductionThere is a growing awareness in the engineering research community of the potential of solid freeform fabrication to greatly expand the space of manufacturable devices by enabling greater geometric freedom in designs and controlled spatial heterogeneity in material composition (Cooper, Kang et al. 1999; Duty, Jean et al. 1999; Li, Golnas et al. 2000; Dybala, Chlebus et al. 2002; Fuller, Wilhelm et al. 2002; Noecker and DuPont 2002; Ouyang, Mei et al. 2002; Stampfl, Hao-Chin et al.

It is our opinion that the majority of applications for this technology in the near term are in the biomedical device and biomimetic robotics fields, with millimeter to decimeter scale. For this reason, we have chosen performance goals for functional components that are similar to the performance of biological functional analogs and more likely to be biocompatible – namely compliant structures and mechanisms, large strain actuation, and low voltage electronics. In prior work, we have demonstrated a multiple material freeform fabrication platform (Figure 3a), and fabrication of batteries, structures, joints and wiring (Malone, Rasa et al. 2004), and more recently, living tissue implants (Cohen, Malone et al. 2004).

2004), and more recently, living tissue implants (Cohen, Malone et al. 2004). The focus of our research has shifted to the fabrication of actuators which are compatible with the aforementioned library and suitable for biomedical devices and biomimetic robotics.484In prior work (Malone and Lipson 2004), we demonstrated that two types of soft electromechanically active materials – conducting polymer (CP) and ionomeric polymer-metal composite (IPMC) – plausibly can be freeform fabricated. We have selected IPMC materials as the most promising on the basis of electromechanical performance. To motivate this decision, a brief review of the mechanisms of CP and IPMC actuation is provided below.BackgroundA typical conducting polymer actuator is a thin polymer film surrounded by a liquid electrolyte (salt solution with mobile ions); an electric potential is applied between the polymer film and an electrode immersed in the electrolyte.Figure 1.

Illustration of the volume change of CP caused by ion diffusion. A CP film doped with large, immobile anions is immersed in electrolyte, and its electrical potential is varied with respect to the electrolyte. During oxidation, the polymer loses electrons and repels small cations in electrolyte. Large anions cannot diffuse quickly into or out of the polymer. During reduction, the polymer gains electrons; small cations can quickly diffuse into the polymer, occupying space and increasing the polymer volume.The CP film actuates by changing volume as ions are forced by the potential to diffuse into or out of the bulk polymer (Figure 1). Conducting polymers are linear chain organic molecules in which successive carbon atoms along the axis of the chain are bound alternately by one shared electron pair (single, or σ-bond) and by two shared electron pairs (double, or π-bond).

By adding or removing electrons from the CP chains, and embedding large positive or negative ions (known as “doping”), respectively, into the chains to maintain electrical neutrality, it is possible to make the locations of the π-bonds less well defined and to vary the electrical conductivity of the polymer (MacDiarmid 2001). With an electric potential between the CP and the electrolyte, free ions (and attached solvent molecules) from the electrolyte diffuse into or out485from between the tangled polymer chains to compensate for the changed charge of the polymer, thereby causing the overall volume of the polymer to change (Figure 1) (Bar-Cohen 2001).An Ionomeric Polymer-Metal Composite (IPMC) actuator consists of a polymer film whose surfaces are partially penetrated by conductive (typically metal) particles. The surfaces are also typically chemically plated or electroplated with metal to increase the surface conductivity.

The surfaces are also typically chemically plated or electroplated with metal to increase the surface conductivity. IPMC’s actuate by bending in response to an electric field applied via these conductive surface electrodes. The electromechanically active polymer is what is known as an “ionomer” - a polymer which has ionic termination, typically on a side branch. Several commercially produced ionomers are available (e.g. Nafion from DuPont Inc., Flemion from Asahi Glass). Nafion is a modified PTFE (Teflon) with perfluorinated sulfonate anion side branches, and is typically produced in what is called “acid form,” meaning that a proton is associated counter-ion to the sulfonate anions in the polymer side chains (Figure 2a). The standard approach to IPMC fabrication (Kim and Shahinpoor 2003) involves purchasing a solid membrane, and replacing the proton by another cation (e.g. Li+) to improve actuation properties (Figure 2a).(a)(b)Figure 2.

Li+) to improve actuation properties (Figure 2a).(a)(b)Figure 2. (a) Structure of DuPont Nafion polymer; schematic exchange of cation (from H+ to Li+) to improve actuation; Rf is proprietary fluorinated group, (b) possible IPMC actuation mechanism: pressure differential from electric field driven ion and solvent motion (adapted from (Shahinpoor and Kim 2001), Figure 8).The surfaces of the membrane are given metallic electrodes, typically by soaking in a platinum salt solution, then chemical reduction of the platinum salt to yield platinum nanoparticles in the outer few micrometers of the membrane. Platinum is the preferred metal because it is immune to corrosion over a larger range of electrochemical potential than other metals, and hence allows the use of higher driving voltages without damage.

The isolated nanoparticles It appears to be essential to the actuation of the IPMC that these conductive particles be dispersed through a finite depth of the membrane in order to increase the depth of the electrostatic double-layers, and hence the fraction of the material which experiences the electric field (Nemat-Nasser and Li 2002). The platinum reduction is then often followed by a surface chemical- or electro-plating, typically with gold, to reduce the surface resistance of the electrodes. The bending of an IPMC in response to an applied electric field seems to result from a combination of mechanisms. These486include the field-driven diffusion of cations and associated solvent (the anionic side groups on the polymer being immobile) causing an internal pressure gradient (Figure 2b), a change in electrostatic forces between the anionic side groups, and reorientation of the (polar) solvent molecules (Shahinpoor and Kim 2001; Nemat-Nasser and Li 2002).

If the actuation signal is rapidly applied (e.g. a step), there is a rapid bending toward the negative electrode, followed by a slower relaxation back toward the positive electrode, resulting in a net motion toward the positive electrode. The fast processes are probably cation and solvent motion, electrostatic stiffness changes, polymer conformation, and solvent orientation related, while the slow process is probably solvent diffusing backward to equalize pressure through the material (Nemat-Nasser and Li 2002).The differences in actuation mechanism for the two types of material have significant practical implications. One of the primary drawbacks of the mechanism of CP actuation for macroscopic actuation is the fact that the polymer changes volume only in proportion the actual volume of ions diffused into or out of the material (Baughman 1996; Benslimane, Gravesen et al. 2000; Bar-Cohen 2001; MacDiarmid 2001; Madden, Madden et al. 2001; Smela and Gadegaard 2001).

2004). Thus, at 1Hz operation, a 5cm by 2cm by 100µm CP actuator (10-7 m3) requires >1A current to actuate, vs. 100mA or for an IPMC of same size. In mobile robotics and other power constrained applications this results in a very significant difference in the scale of the power storage or generation required, and the large thermal dissipation is undesirable in biomedical applications. Because a CP actuator is a capacitive structure, the electromechanical efficiency of CP actuators can be improved by recovering energy stored, but this requires much effort to reduce resistive losses in CP and associated electronics, and/or very slow actuator operation to be useful at the macroscopic scale. In any case, the electromechanical efficiency of CP actuators is in the vicinity of 0.1% without energy recovery, at least for polypyrrole, and could conceivably reach ~10% with perfect recovery and a restricted regime of actuation (Madden, Madden et al. 2001).

2001). By comparison, IPMC actuators are routinely achieving 1.5% efficiency (Shahinpoor and Kim 2001; Kim and Shahinpoor 2003), with significant gains possible through more careful control (Jung, Nam et al. 2003). IPMC devices are also capacitive, and subject to similar efficiency improvements via stored energy recovery. Given also that the achievable strains and stresses are roughly comparable for the two types (Madden, Vandesteeg et al. 2004), and that IPMC actuators can generate large scale motion at up to 100Hz (Nemat-Nasser and Wu 2003; Paquette and Kim 2004) in air, IPMC actuators are superior for rapid, power constrained, macroscopic actuation.The main drawbacks to IPMC actuators, by comparison to CP, are a lack of a “no power catch-state” (Madden, Vandesteeg et al.

2004), in which the device maintains a given actuated strain with essentially no power input, and the fact that IPMC’s actuate strictly by bending, without volume change (Shahinpoor and Kim 2001; Nemat-Nasser and Li 2002). This last issue implies that adapting IPMC to linear actuation requires complex geometry or a mechanism of487some sort. Certainly many applications exist for bending mode actuation (Jung, Tak et al. 2003; Kim, Ryu et al. 2003; Paquette and Kim 2004), and some commercialization of IPMC is underway, with biomedical devices such as heart-contraction assistance devices (Shahinpoor 2002), vision correction devices (Shahinpoor, Shahinpoor et al. 2003), intraocular pressure management devices for glaucoma management (Soltanpour 2001), a variety of biomedical micropumps (Soltanpour and Shahinpoor 2003; Soltanpour and Shahinpoor 2004), and toys (Ashley 2003) currently the target markets.

The utility of linear actuation is difficult to argue against for most applications however, and the problem of manufacturing linearly actuating IPMC devices is starting to be addressed in the research literature (Kaneda, Kamamichi et al. 2003; Yamakita, Kamamichi et al. 2004). We believe that the geometric freedom and spatial control over material placement provided by freeform fabrication will enable the production of linear IPMC actuators that exceed the performance of the state of the art.ExperimentalThe traditional approach to manufacturing IPMC described above is not readily adapted to freeform fabrication in that a prefabricated solid ionomer membrane must be chemically treated and metallized in several separate baths. Kim and Shahinpoor (Kim and Shahinpoor 2001) present a method of producing IPMC devices which involves Nafion ionomer membranes made by casting a liquid dispersion (“Liquid Nafion”).

Kim and Shahinpoor (Kim and Shahinpoor 2001) present a method of producing IPMC devices which involves Nafion ionomer membranes made by casting a liquid dispersion (“Liquid Nafion”). The authors note that the cast dispersion is very prone to cracking unless an additive (typically a low vapor pressure solvent) is added to the dispersion and solvent evaporation is gently performed. In addition, the resulting membrane must be annealed at elevated temperature in order to develop some crystallinity. Without annealing, the ionomer would remain soluble, and actuation (which requires the membrane to be saturated with solvent) would not be possible. This “solution recasting” method permits the thickness of the membrane to be customized to the actuator application, and the authors present a 2mm thick IPMC which demonstrates the significant output force increase which is possible. Note that this device is still given surface electrodes via metal reduction and plating methods.

Note that this device is still given surface electrodes via metal reduction and plating methods. The authors suggest (but do not report results from) a method of producing the electrodes and the membrane entirely by casting stacked layers, where the electrode layers would be cast from an ionomer dispersion with suspended metal particles. This method would permit freeform fabrication of complete IPMC devices, and was taken as the starting point for the present efforts.Initial Formulation and Hand-Made IPMCWe performed an initial set of experiments which was intended to culminate in the production of a working IPMC by hand via the stacked casting method described above. 5wt% H+ Nafion dispersion in mixed alcohol and water (Sigma Aldrich prod. # 527084) was used as the ionomer, and an initial electrode material consisted of silver powder (99.9%, 1-3µm particles) mixed in the ionomer dispersion. Silver is used in these experiments as a lower cost alternative to platinum.

Silver is used in these experiments as a lower cost alternative to platinum. An annealing schedule of 45 minutes at 70C in air was found to eliminate solubility of the ionomer without causing brittleness from too high a degree of crystallinity. Casting the neat ionomer alone atop a PTFE substrate revealed that good films of at least 250um can be formed without cracking and without the addition of any additives. When cast atop a dried film of the electrode material, however, two problems occurred: firstly, the dried ionomer in the electrode layer is highly hydrophobic, and causes subsequently deposited ionomer to migrate, bead up, and crack upon drying, and secondly, the dried electrode layer (whether annealed or not)488absorbs solvent from the newly applied ionomer, causing swelling, buckling and cracking of both materials.

It was found that the addition of a non-ionic surfactant (DuPont Zonyl FSO-100) in a low concentration (~ 2 wt%) to the electrode material increased the affinity of the electrode material sufficiently to remove the first problem. The second problem can be minimized by deposition of subsequent layers atop prior layers that have not completely dried. This complicates the deposition of the upper electrode layer in that the suspended silver particles tend to settle during drying, and can penetrate the neat Nafion layer, causing internal electrical shorting of the finished IPMC. Careful timing and control of the deposited quantity of electrode material reduces but does not eliminate this problem.We made several 1cm by 3cm by ~0.250um IPMC actuators by manual deposition onto a PTFE substrate and annealing as mentioned above. The devices were soaked in deionized water, for 30 minutes, and then peeled from the substrate with tweezers.

The devices were soaked in deionized water, for 30 minutes, and then peeled from the substrate with tweezers. The actuation of these in air was verified qualitatively under 1.5V stimulus by videography, confirming that the deposited electrode concept suggested by Kim and Shahinpoor (Kim and Shahinpoor 2001), is at least feasible.Unfortunately, the devices are limited to only a few actuation cycles in air because of solvent loss through the surfaces and progressive development of internal electrical shorting – perhaps through migration of silver particles, or because the internal ionomer layer is too thin or flawed. In addition, the materials as formulated above, and the process described are not completely satisfactory for automated freeform fabrication. The silver particles in the electrode material settle out of suspension very quickly making the material difficult to dispense.

The silver particles in the electrode material settle out of suspension very quickly making the material difficult to dispense. Some means of confining the deposited liquids is required to allow freeform fabrication of an IPMC in an arbitrary two dimensional shape, and ideally this confinement should allow the actuator to be fabricated as an embedded component of a larger freeform fabricated device.Containment, and Reformulation for Freeform FabricationOur second series of experiments focused on developing a means of freeform fabricating a container into which the IPMC materials can be cast, and on modifying the electrode material formulation to improve dispensability. One of the materials that we have successfully employed for freeform fabrication is poly(dimethylsiloxane) or PDMS, also known as silicone.

One of the materials that we have successfully employed for freeform fabrication is poly(dimethylsiloxane) or PDMS, also known as silicone. This material is very chemically inert, thermally stable to over 200C, we can routinely achieve 500µm resolution when fabricating structures with it, and can deposit it with good control on most of the other materials that our system employs. Rectangular silicone wells (Figure 3b), with 1cm by 3cm by 2-5mm internal dimensions were deposited onto glass microscope slides for use as containers for IPMC fabrication. We cast several actuators into these by hand to verify their compatibility with the existing materials and process. We observed that the neat ionomer dispersion has a higher affinity for the silicone than for the electrode material, and there is a tendency for the ionomer to seep under the lower electrode, and also to be drawn toward the boundaries of the well. This apparently produces stresses in the film as it dries, leading to cracking.

This apparently produces stresses in the film as it dries, leading to cracking. Adding more surfactant to the electrode material exacerbates the seepage problem such that the electrode becomes surrounded by ionomer, and electrically insulated. We added a low-vapor pressure solvent (N-N-dimethylformamide, DMF) to the ionomer dispersion to help control cracking. At 33wt% of DMF and above, the cracking problem was eliminated, at the cost of greatly increased drying time. We modified the electrode material to be more readily489dispensed by mixing in a quantity of carbon black powder along with the surfactant and silver powder in the ionomer dispersion.(a)(b)Figure 3.

(a) Schematic of multiple material freeform fabrication system with Cartesian gantry positioning system and multiple changeable deposition tools, (b) a freeform fabricated silicone well.This results in a stable slurry, referred to as Electrode Material 6 (EM6), which can be dispensed from a 0.5mm ID syringe needle without clogging. A four-probe measurement technique was used to obtain the resistivity of this material; the values for pure silver, carbon, and platinum are provided for reference.Table I.

A four-probe measurement technique was used to obtain the resistivity of this material; the values for pure silver, carbon, and platinum are provided for reference.Table I. Formulation and Resistivity of Electrode Material 6Material Ag Powder, 99.9%, 1-3um H+ 5wt% Nafion Dispersion DuPont Zonyl FSO-100 Surfactant Carbon blackValueUnits 20 wt% 70 wt% 2 wt% 8 wt%Resistivity EM6 (3 samples, 4-probe method) Silver, pure solid (Matweb 2005) Carbon, graphite (Matweb 2005) Platinum, pure solid (Matweb 2005)Units 3.85(± 3.76) E-03 Ω cm 1.77E-06 Ω cm 6E-03 Ω cm 1.06E-05 Ω cmValueIn order to be able to deposit liquid materials precisely into the silicone well, we extended the manufacturing planning software of our fabrication system to allow structurally constrained prioritization of deposition, which we term “backfill deposition”. As parts of a multiple material assembly are imported into the software, they can be assigned a sequential priority for fabrication.

As parts of a multiple material assembly are imported into the software, they can be assigned a sequential priority for fabrication. This priority will be obeyed by the system except where doing so would violate the relationship of one part supporting another. Thus the silicone well can be given a higher priority than the materials to be deposited into it, and it will be completely fabricated to its full height490before the deposition of the first electrode layer begins (Figure 4a). This greatly improves the quality of cast materials by allowing the deposition nozzle to be at its optimal height above the layer beneath.We deposited the reformulated materials with the freeform fabrication system to produceworking IPMC devices in the following sequence:1. A silicone well is deposited onto a glass slide 2. A layer of electrode material is backfilled into the well and allowed to dry until the surface begins to appear matte.3.

A silicone well is deposited onto a glass slide 2. A layer of electrode material is backfilled into the well and allowed to dry until the surface begins to appear matte.3. The well is backfilled with ionomer dispersion (with DMF additive) and allowed to dry until the surface does not flow.4. The upper electrode material is deposited and allowed to air dry 5. The glass slide is placed in an oven and annealed at 70C for 45minutes 6. The silicone well is filled with deionized water for 30 minutes to saturate the IPMC7. The device is lifted from the well with tweezers and testedThe actuation of these devices was verified qualitatively by videography as before (Figure 4). Once again, the devices suffer from internal shorting, and have a very short service life in air– typically only 2 or 3 actuation cycles. Nevertheless, these are the first completely freeform fabricated IPMC actuators.(a)(b)Figure 4. (a) “Backfill deposition” of electrode layer into silicone well.

Nevertheless, these are the first completely freeform fabricated IPMC actuators.(a)(b)Figure 4. (a) “Backfill deposition” of electrode layer into silicone well. (b) Two superposed frames from video of actuation test of freeform fabricated IPMC. In air, 1.5V step, elapsed time 5s.Reformulation for Improved Lifetime and EfficiencyIn a third sequence of experiments, we have focused on improving the service life and reducing the input power requirements (internal shorting) of freeform fabricated IPMC actuators. It is our experience that the primary limitations on the service life of IPMC in air are the loss of491solvent, corrosion of the electrodes, and progressive electrical shorting. As described above, the IPMC actuation mechanism involves, in part, the generation of an internal pressure gradient by the electrophoretic transport of ions and solvent. This pressure gradient drives solvent out through the porous electrodes.

This pressure gradient drives solvent out through the porous electrodes. Given time it can be reabsorbed into the actuator (Nemat-Nasser and Li 2002), but while on the surface it is likely to be lost to evaporation, dripping, or contact transfer. Internal shorting causes resistive Joule heating which increases solvent evaporation. It has also been noted (Kim and Shahinpoor 2003) that solvent leakage probably also reduces the maximum achievable force output from the actuator by limiting the internal pressure gradient that can be maintained.In order to reduce the rate of solvent loss, we have explored the use of impermeable electrode materials which can trap the solvent inside of the IPMC.

There are several essential requirements that such a material must meet:• • • • •Low modulus of elasticity Low electrical resistivity Solvent (water) impermeability Chemically inert to other IPMC materials Good adhesion to other IPMC materials Easily depositedThe impermeable material is intended to serve as a surface coating to the silver-filled ionomer electrode layers. As mentioned above, some conductive particle penetration into the membrane is essential to the actuation of IPMC.We performed a set of experiments involving 13 candidate materials in order to arrive at a material which satisfactorily meets all of these requirements. The concept we employed in developing these materials was to form a composite of metal particles embedded in a hydrophobic and/or impermeable elastomer matrix. The matrix provides the solvent barrier and low modulus, while the metal particles reduce electrical resistivity of the polymer.

The matrix provides the solvent barrier and low modulus, while the metal particles reduce electrical resistivity of the polymer. If the elastomer is a thermoplastic, then dispensability is achieved by dissolving the material in an organic solvent which is compatible with the other IPMC materials, or the material could be melted and extruded. If it is thermosetting, then one-part, room-temperature vulcanizing (RTV) materials are preferred so that elevated temperature curing and crosslinking agents are not required.The 13 materials explored employed one of two elastomers - silicone and “Hydrin”. The majority of metal-filled silicone formulations that we explored failed to achieve reasonable electrical resistivity, and were rejected. A silver-filled RTV silicone (SS-26, Silicone Solutions Inc.) has acceptable resistivity.

A silver-filled RTV silicone (SS-26, Silicone Solutions Inc.) has acceptable resistivity. Unfortunately, it is stiff from the high concentration of metal particles, and does not adhere well to the other IPMC materials - IPMC’s produced with it delaminate immediately.We found Hydrin C thermoplastic (poly(epichlorohydrin-co-ethylene oxide), Zeon Chemicals L.P.) to be a reasonably good elastomer for this application. Low resistivity is achieved without dramatic increase in stiffness, it is hydrophobic but soluble in several common492solvents including toluene, acetone, and DMF, and it adheres well to the IPMC materials. The main drawback to the material is a somewhat higher permeability to water vapor than some alternatives (Table II).Table II.

The main drawback to the material is a somewhat higher permeability to water vapor than some alternatives (Table II).Table II. Properties of some candidate elastomer materialsElastomer Hydrin C ("Epichlorohydrin Elastomers: Hydrin ECO" 2005) Butyl Rubber (Gauthier 1995) EPDM Rubber (Gauthier 1995)Water Vapor Transmission Rate at 38C (g/(m2 * 24h))85.870.2176.045Tensile Modulus, 25C, 100% Strain (MPa)3.10.3-3.40.7-20.7Tensile Strength (MPa)13.213.83.4-24.1Shore A Hardness8530-10030-90Suspending 1-3µm silver particles in dissolved Hydrin is difficult. A commercial “flexible silver ink” (EP403920-50, SRA Inc.), which uses specialty polymers to suspend the particles without adversely affecting the conductivity was found to form a good alloy with Hydrin. Both materials are soluble in toluene, which evaporates quickly and does not attack the ionomer.

Both materials are soluble in toluene, which evaporates quickly and does not attack the ionomer. The addition of surfactant was again necessary to preventing migration and cracking of subsequent layers cast atop this material. Table III presents our best formulation, referred to as Electrode Material 19 (EM19), its resistivity and that of some reference materials.Table III. Formulation and Resistivity of Electrode Material 19Material Zeon Chemicals L.P. Hydrin C Elastomer Toluene SRA Inc. EP403920-50 Flexible Silver Ink DuPont Zonyl FSO-100 SurfactantValueUnits 6 wt% 72 wt% 20 wt% 2 wt%Resistivity EM19 (3 samples, 4-probe method) Silver, pure solid (Matweb 2005) SRA Inc. EP403920-50 Flexible Silver Ink (Stan- Rubenstein-Associates 2005) Silicone Solutions Inc. SS-26 Silver-filled RTV (Silicone- Solutions-Inc.

EP403920-50 Flexible Silver Ink (Stan- Rubenstein-Associates 2005) Silicone Solutions Inc. SS-26 Silver-filled RTV (Silicone- Solutions-Inc. 2005)Units 6.96(± 0.96) E-03 Ω cm 1.77E-06 Ω cmValue1E-04 Ω cm5E-03 Ω cmWe have dealt with the internal shorting problems by increasing the thickness of the neat ionomer layer. Thus a larger volume of liquid Nafion must be deposited and the drying time (especially given the low vapor pressure of the DMF additive) is very long. We employ a more concentrated ionomer dispersion to mitigate this. Switching to 20wt% H+ Nafion dispersion (Sigma-Aldrich prod. # 527122) requires increasing the amount of DMF added up to 50% by weight to prevent cracking, but still results in greatly reduced time for evaporation. We modified the freeform fabrication process to employ these materials:4931. A silicone well is deposited onto a glass slide. 2. A layer of EM19 electrode material is deposited into the well and allowed to dry completely.3.

A silicone well is deposited onto a glass slide. 2. A layer of EM19 electrode material is deposited into the well and allowed to dry completely.3. A layer of EM6 electrode material is deposited into the well and allowed to dry until matte in appearance.4. The well is backfilled to the top with ionomer dispersion (with DMF additive) and allowed to dry until the surface does not flow.5. Repeat step 3. 6. Repeat step 2. 7. The glass slide is placed in an oven and annealed at 70C for 45minutes. 8. The silicone well is filled with deionized water for 30 minutes to saturate the IPMC.9. The device is lifted from the well with tweezers and tested.Figure 5 shows the CAD model of the devices that we produced with our fabrication system, and the appearance of a finished device.(a)(b)Figure 5.

Freeform fabricated 5-layer IPMC, (a) cutaway of CAD model showing layer sequence, (b) IPMC in its well (top), and after hydration and removal (bottom).ApparatusTo provide more quantitative testing of the performance of the IPMC devices being produced, we developed a test apparatus (Figure 6) which interfaces with a PC-based data acquisition system (IOTech DAQBoard 2000, running DASYLab DAQ software). This apparatus allows the application of a controlled voltage waveform to the actuator, while simultaneously collecting applied voltage, current, and force output by the actuator. The actuator is clamped on one end to a 12g counterweight which rests atop a laboratory scale with a 5mg resolution (ADAM ACB Plus-150). The other end of the actuator is held in a spring clamp, the tips of which form isolated electrical contacts. The spring clamp is held in a laboratory ring stand.

The other end of the actuator is held in a spring clamp, the tips of which form isolated electrical contacts. The spring clamp is held in a laboratory ring stand. A power operational amplifier receives a voltage command from the DAQ system, and attempts to maintain this voltage across the actuator electrodes by applying a current to the494actuator via the spring clamp contacts. The applied current is sensed by the data acquisition system by analog to digital (A/D) conversion of the voltage drop across a current sensing resistor in the current path. The actual voltage achieved by the amplifier is also measured by A/D at the output of the amplifier. The force output is sensed by the scale, which streams readings via serial communications at about 40Hz to the DAQ system.Power ClampCounterweight ClampApplied VoltagePower Amp+RsenseIPMC Actuatore c r o FData Acquisition SystemMeasured CurrentMeasured VoltageSCALESTANDScale RS-232Figure 6.

Schematic of test apparatus for acquiring force output, voltage, and current versus time.The life, and correlate applied signals with force output.simultaneous collection of these data allows us to calculate applied power, measure cycleResults ave tested several 5-layer freeform fabricated devices using the above apparatus. We h s a large variation in the output force from cycle to cycle, and the power delivered to the There i devices is very noisy (see Figure 7). This latter effect is probably a result of the particulate nature of the conductive phases of the electrode materials making and breaking conductive paths as the device bends and solvent migrates. A discrete Fourier transform of the input and output signals clearly demonstrates correlation. At this time, all testing takes place with a 0.1Hz, 1.5V square wave applied signal, and the output dynamic force can be seen to peak strongly at this frequency and its harmonics (Figure 8).

At this time, all testing takes place with a 0.1Hz, 1.5V square wave applied signal, and the output dynamic force can be seen to peak strongly at this frequency and its harmonics (Figure 8). In order to extract the maximum blocked force for each cycle in the large datasets, peaks in the force data are automatically identified using the zeros of the first difference of the force data. Figure 9 provides a detail of the results of this algorithm. Our highest performance device achieved more than 250 (bi-directional actuation, 513 total peaks) cycles in air, operating continuously without rehydration for about 40 minutes.Its max um blocked force was 0.678mN, and average of peak blocking force over the entire testing period was 0.150 (+/-0.130)mN. The RMS power delivered to the actuator was 9.4mW. Table IV presents these results in comparison to some of the best results published for IPMC actuators produced by any means.im495Input Power and Dynamic Force vs.

Table IV presents these results in comparison to some of the best results published for IPMC actuators produced by any means.im495Input Power and Dynamic Force vs. Time0.110.080.80.060.60.040.4]N m]W[r e w o P0.0200.020.200.2[ e c r o F c m a n y Di0.040.40.060.60.080.80.1100 100110 110120 130 130 120 Elapsed Time [s]140 1401 150 150Figure 7. Detail of force and consumed power vs. time for best freeform fabricated IPMC100Normalized Spectral Power102r e w o P d e z104i la m r o N10610810100Input Voltage Input Power Output Dynamic Force0.050.10.150.2 Frequency [Hz]0.250.30.350.4Figure 8.

Comparison of performance of freeform fabricated IPMC to other published IMPC resultsTypeDimensionsTest ConditionsForce (mN)Est. Shear Stress (MPa)Est. Shear Stress/Power (MPa/W)1Freeform fabricated H+ Nafion IPMC Recast Li+ Nafion IPMC, standard Pt/Au(Kim and Shahinpoor 2002) Standard Li+ Nafion IPMC, standard Pt/Au (Shahinpoor and Kim 2001) Stretched Li+ Nafion IPMC, standard Pt/Au (Kim and Shahinpoor 2003)20mm by 9mm by 1mm15mm by 5mm by 2mm20mm by 5mm by 0.2mm10mm by 5mm by 0.2mm1.5V, 0.1Hz square 0.6782V, 0.5 Hz square601V, 0.5Hz square161.5V, 0.5Hz square601.5E-034.5E-021.63.01.6E-019.0E-016.4E+018.0E+01DiscussionAs presented in Table IV, the best reported blocked forces (fourth row) are nearly 100 times higher than those of our freeform fabricated device.

Scaling to shear stress per unit power increases this gap to a humbling 500 times, but interestingly, the “recast” IPMC (second row) is merely 5.62 times better than our freeform fabricated IPMC in this measurement. The recast IPMC was made by casting the ionomer membrane from liquid dispersion, followed by traditional platinum and gold electroding (Kim and Shahinpoor 2002). The similarly poor performance of the devices made with liquid ionomer dispersions suggests that some microstructural traits of a cast membrane make it inferior (for actuation) to the commercially produced ones. Kim and Shahinpoor (Kim and Shahinpoor 2002) report that passive tensile stress/strain behavior is different for cast and commercial membranes. DuPont produces Nafion membranes both by casting and by extrusion – unfortunately the type used is never mentioned in the literature.

DuPont produces Nafion membranes both by casting and by extrusion – unfortunately the type used is never mentioned in the literature. It should also be noted that the freeform fabricated devices are using H+ as the counter-ion in the ionomer for simplicity; it is well documented (Kim and Shahinpoor 2003) that use of other counter-ions can improve blocked force by 2-3 times, suggesting that the remaining deficit may be accounted thusly.Future WorkAs mentioned above, all of the above experiments have, for simplicity, made use of Nafion ionomer dispersion as purchased, in what is referred to as the “acid form” - that is with a proton as the counter ion which is associated with the anionic side chains of the ionomer (Figure 2a). As has been observed in the literature (Nemat-Nasser and Li 2002; Kim and Shahinpoor 2003), the force output of IPMC actuators is dramatically affected by the choice of counter ion, with Li+ offering perhaps more than 2.5-fold increase in force over H+.

We have already tested a convenient method of converting the counter ion to Li+ by the simple addition of LiOH to the ionomer dispersion, and are in the process of investigating the effects of this change on the other1 The referenced literature does not contain simultaneous data for force and power, nor for force, voltage and current. We have used selected 25mA as a reasonable average current based on Kim and Shahinpoor 2003, Figures 11-13.498materials and the manufacturing process. Several other modifications may also yield large improvements in performance, including reducing the size of conductive particles used in Electrode Material 6 to the nanometer level (Kim and Shahinpoor 2001), forming a nanocomposite of the ionomer with Montmorillionite clay particles (Paquette, Kim et al. 2003), and using a lower modulus, more vapor impermeable elastomer in Electrode Material 19.

2003), and using a lower modulus, more vapor impermeable elastomer in Electrode Material 19. Achieving truly three dimensional actuators will require modifying the materials to achieve very rapid drying and thin layers out of the horizontal plane. Figure 11 presents a concept for a three dimensional actuator geometry which would produce linear actuation.(a)(b)Figure 11. Concept of 3-dimensional IPMC geometry which converts bending motion to linear motion, (a) relaxed, (b) contracted.It is our aim to approach the performance of conventionally produced IPMC’s, and then to exploit the geometric control afforded by freeform fabrication to develop IPMC designs which are customized for applications and integrated with devices to a degree not otherwise possible.ConclusionIn the quest to develop a complete library of functional electromechanical elements which can be freeform fabricated and are mutually compatible, we have been investigating electromechanically active materials.

Ionomeric Polymer-Metal Composite materials have been selected as the most compatible with our fabrication system, with the freeform fabricated functional elements that we already produce, and with the goal of producing actuation suitable for biomedical devices and bio-inspired robotics. We have developed new material formulations and manufacturing processes which have allowed us to produce the first ever completely freeform fabricated IPMC actuators. The performance of these devices has been quantitatively measured. We have presented the first cycle life measurements for in-air operation of an IPMC. The output blocked force measured is inferior to that produced by IPMC devices produced in the standard manner, but we are in the process of incorporating materials modifications (e.g. conversion of the ionomer to Li+ form) which are well-known to produce multifold increases in force output.

conversion of the ionomer to Li+ form) which are well-known to produce multifold increases in force output. The freeform fabrication process developed includes the fabrication of wells to contain the liquid active materials during their casting. In that these wells can be deposited at an499arbitrary location, atop almost any other material, this represents a major step toward fabricating actuators as integral components of complete, freeform fabricated electromechanical devices.Acknowledgments This work was supported in parts by the U.S. Department of Energy, grant DE-FG02- 01ER45902.References Ashley, S. (2003). "Artificial Muscles: Novel motion-producing devices--actuators, motors,generators--based on polymers that change shape when stimulated electrically are nearing commercialization." Scientific American 289(4): 53-61.Bak, D. (2003).

tentials of tens of volts or higher to switch, requiring more printed battery cells to power them than are feasible to fabricate. The organic electrochemical transistor, or OECT (White et al., 1984), is a more feasible option that we consider here. Unlike OFETs that rely on electrostatic gating, OECTs rely on electrochemical doping/de-doping to switch the transistor on and off. These devices have low (< 1V) operating voltages, their performance does not depend heavily on their dimensions, and the entire transistor can be fabricated under ambient conditions.We have successfully produced functional OECTs using one of our multi-material freeform fabrication systems. Their electrical performance, which we have characterized, is qualitatively similar to the reported performance of similar OECTs in the literature (Nilsson et al., 2005).

Their electrical performance, which we have characterized, is qualitatively similar to the reported performance of similar OECTs in the literature (Nilsson et al., 2005). Having produced these working devices on glass substrates, we have identified several printable substrate materials for the next generation of devices, and are in the process of characterizing them. Using the information gained from these efforts, we have added inkjet deposition capability to one of our SFF systems, improving the machine’s ability to produce high-resolution thin films.

Using the information gained from these efforts, we have added inkjet deposition capability to one of our SFF systems, improving the machine’s ability to produce high-resolution thin films. These developments will enable us to produce more compact transistors with higher performance and better reproducibility, which can be integrated into larger fully-printed electromechanical devices.Background Most transistors used in electronic devices today are fabricated on single-crystal silicon wafers, have micron and sub-micron sized features, and require large and very costly equipment in clean- room environments for processing. The most promising materials for printable transistors are organic semiconductors, which are often solution-processable and do not demand such stringent environmental constraints during processing.

The most promising materials for printable transistors are organic semiconductors, which are often solution-processable and do not demand such stringent environmental constraints during processing. While spin-coating and photolithography are used to define the features of many organic transistors (Ziemelis, 1998), transistors and even entire circuits fully printed with ink jet or screen printing mechanisms have also been developed (Paul et al., 2003; Katz, 2004; Berggren et al., 2007). These advances have opened up the possibility of printing working transistors on a low cost freeform fabrication system, or “fabber,” which can also print the wires, batteries, actuators, and structural components required to fabricate an entire working electromechanical device.Working with such a versatile system, however, means handling various other constraints.

All materials must be “addition-processable” – this usually implies deposition of liquids, but may also include vapor deposition, as with Optomec’s M3D process, (Hedges et al., 2005) or controlled solidification (as in stereolithography). This requirement includes the active materials in the transistor as well as electrically conductive materials used for electrical contacts and wires. We also need to develop a printable, electrically insulating substrate for our transistors, allowing them to be integrated into larger fully-printed devices. All of these materials should be processable in ambient conditions; we can anneal while printing with our research platform, but many materials require treatment in an oxygen-free environment, or release noxious fumes during drying or annealing. Because of the physical limitations of our current syringe-based deposition process, the transistors cannot have any lateral dimensions smaller than 250 micrometers.

Because of the physical limitations of our current syringe-based deposition process, the transistors cannot have any lateral dimensions smaller than 250 micrometers. A drop-cast or printed film of material, in theory, can be made arbitrarily thin, as a very thin material could be produced by printing that material from a very dilute solution, but it is difficult to ensure the robustness and uniformity of such a film. Finally, if we wish to incorporate these transistors into fully-printed electromechanical devices, the transistors need to run at voltages and currents low enough that the transistors can be powered with a reasonable number of printable battery cells.We are in the process of adding an IJP (inkjet printing) head to a Fab@Home Model 1 freeform fabrication system (Malone and Lipson, 2007).

The IJP head is compatible with at least two candidate materials for producing transistors, and should provide many advantages over syringe deposition, the default method of our systems, especially in terms of the smallest printable lateral dimensions and the ability to produce thin films. The constraints we need to overcome are still similar to those noted above, however. The IJP head is also limited to very low viscosity and carefully filtered materials, so the fabrication of transistors embedded within other structures will require a combination of IJP and syringe deposition methods.One very common type of transistor is the field-effect transistor (FET). The FET category encompasses the type of silicon-based transistors used in most electronic devices today, but FETs have also been made with different combinations of organic and inorganic materials (Horowitz, 1998; Dimitrakopoulos and Mascaro, 2001), and have been produced by inkjet printing (Paul et al., 2003; Katz, 2004).

However, because FETs operate under the local effects of electrostatic fields, very small changes in the dimensions of a FET can have significant effects on its performance. Specifically, fabricating FETs necessitates printing very thin, but robust, insulating films, which are difficult to produce reliably on the required scale with any of our setups. FETs often require tens of volts to switch on and off (for example, see Horowitz et al., 1998) as well. For these reasons, they would be difficult to fabricate with our current freeform fabrication system.Another type of transistor, known as an organic electrochemical transistor (OECT), has been developed (White et al., 1984), and shows more promise for compatability with freeform fabrication.

These transistors, although not as common as FETs, have been used for applications in biological and chemical sensing (Mabeck and Malliaras, 2006; Bernards et al., 2006), and are appealing for use in printable applications as well.OECTs are composed of an organic semiconductor, an electrolyte, and a gate electrode. One active semiconductor material used to make OECTs is poly(3,4-ethylenedioxythiopene) doped with poly(styrene sulfonate) (PEDOT:PSS – see Groenendaal et al., 2002 for general information), which is a degenerately doped p-type semiconductor. By applying a voltage to the OECT gate terminal, it is possible to modulate the resistance of the active semiconductor material “channel” between the transistor’s source and drain terminals (see Figure 1) by electrochemically doping or de-doping the PEDOT:PSS. The relevant electrochemical reaction is:PEDOT:PSS + M+ + e- (cid:1)(cid:1)(cid:1)(cid:1) PEDOT + M:PSSwhere M is some cation (the case for monovalent charge is shown).

The relevant electrochemical reaction is:PEDOT:PSS + M+ + e- (cid:1)(cid:1)(cid:1)(cid:1) PEDOT + M:PSSwhere M is some cation (the case for monovalent charge is shown). PEDOT:PSS has a substantially greater conductivity than PEDOT. When a positive voltage is applied to the gate with respect to the grounded source, cations from the electrolyte migrate into the channel and de-dope the PEDOT:PSS. As a result, the conductivity of the channel decreases, causing a measurable decrease in the source-drain current. (For more details on OECT operation, see Bernards and Malliaras, forthcoming.) For devices that utilize PEDOT:PSS as the gate electrode, the reverse (oxidation) reaction occurs at the gate to maintain charge balance.

Because PEDOT:PSS can be easily over-oxidized (Nilsson et al., 2005), causing it to lose its conductivity permanently, the gate is made much larger (>10x) in area than the channel to prevent over-oxidization of the PEDOT:PSS gate.GM+M++M+VGM+M+DM+M+M+– SFigure 1 – A schematic of an OECT designed to be printed laterally. The blue material represents an electrochemically-switchable material, which the yellow material represents an electrolyte (containing M+ cations).The resulting I-V behavior from PEDOT:PSS OECTs was published by Nilsson et al. in 2002 and 2005. When a negative potential is applied to the drain, current saturation is observed as the magnitude of the drain potential increases, which is desirable if the transistor is to be used as part of an amplifier. A positive potential applied at the gate decreases the amount of current that flows between the source and drain.

A positive potential applied at the gate decreases the amount of current that flows between the source and drain. The behavior exhibited by transistors biased with negative drain and positive gate potentials looks qualitatively similar to the behavior expected from depletion mode field-effect devices. These OECTs require less than 1V between the gate and source to switch from “on” to “off” as well, meaning that they could be powered with a single printed battery cell.It is not difficult to find mutually insoluble and compatible materials to make OECTs. They can be printed laterally (as in Figure 1); the gate, source, and drain can all be printed simultaneously on a single layer with a single material. Fewer layers mean fewer solvent compatibility issues, as well as a device that is easy to print. PEDOT:PSS is available commercially as an aqueous dispersion, but a dried film of PEDOT:PSS is relatively robust in water.

PEDOT:PSS is available commercially as an aqueous dispersion, but a dried film of PEDOT:PSS is relatively robust in water. This property opens the possibility of using an aqueous solution as the electrolyte material. Additionally, we have previously printed conducting “wires” using commercial colloidal silver ink and also by depositing molten Pb-Sn alloy (Malone et al., 2004); the silver ink is also appropriate for the metallic contacts needed to connect the OECTs to other devices, as it employs a solvent that does not dissolve the PEDOT:PSS. This gives a full set of materials suitable for ECT production that are easy for us to work with.OECTs are also less sensitive to size than FETs. They rely on bulk electrochemical reactions in the active material to switch on and off, making distances less important than is the case with field- effect transistors.

They rely on bulk electrochemical reactions in the active material to switch on and off, making distances less important than is the case with field- effect transistors. None of the printed layers of OECTs need to be extremely thin, and OECTs with key features on the order of millimeters are functional, as we report here and as is reported elsewhere (Robinson et al., 2006). These same properties of OECTs also make them significantly slower than FETs, but this is acceptable for our current needs. Since we are limited in the size and reproducibility of the lateral dimensions and thicknesses of the materials we print, OECTs are a good starting point for freefrom fabricated transistors.PEDOT:PSS-based devices are, additionally, very versatile. PEDOT:PSS transistors can be made to mimic the behavior of either “n-type” or “p-type” FETs (“n” and “p” designate the type of charge carriers in the FET semiconductor; both types are used in logic circuitry) (Chen, 2005).

This is an improvement over OFETs, since organic n-type semiconductors tend to be unstable. PEDOT:PSS-based electrochemical devices have been used to make rectifiers (Chen, 2006), logic circuits, and ring oscillators (Nilsson et al., 2005), and have been used as memory devices (Nilsson et al., 2002). Oxidized PEDOT:PSS is dark blue, while reduced PEDOT is lighter; this property has been harnessed to make simple, monochromatic displays (Andersson et al., 2002). PEDOT:PSS can also be used as a conducting material. These different functionalities of PEDOT:PSS have recently been combined into a single fully-printed device on a paper substrate (Berggren et al., 2007). By mastering the use of PEDOT:PSS for OECTs, we gain the ability to use this material in many other devices and systems.Inkjet printing should be able to produce smaller and better-defined patterns of a material than is possible with our current syringe-deposition setup.

An inkjet printer (IJP) head deposits material by ejecting small droplets of a solution at a given spatial frequency onto a substrate, allowing precise placement of relatively small volumes of these materials. One can pattern films of a material from solution by controlling this IJP head with a freeform fabrication system and appropriate software. The deposited film’s thickness is controlled by adjusting the ratio of the translation rate of the print head over the substrate to the firing rate of the jet, within the constraints imposed by the surface energy of the solution on the substrate. The optimum density of droplets per unit area will, in theory, create the uniform films necessary for printing electrically active devices.IJP has two main advantages over syringe deposition.

The optimum density of droplets per unit area will, in theory, create the uniform films necessary for printing electrically active devices.IJP has two main advantages over syringe deposition. First, such a tiny volume of material can be deposited (picoliters to nanoliters) with such high repeatability that material dries or solidifies very quickly, and lateral positional accuracy is determined almost entirely by the positioning system, rather than by material relaxation or flow. Second, achieving precise control of material flow from a syringe requires the syringe needle remain very close to – but not touching – the substrate, so that the deposited flow does not break irregularly into droplets and the needle does not collide with previously deposited material. This is exceedingly difficult to achieve without sophisticated sensing and feedback control.

This is exceedingly difficult to achieve without sophisticated sensing and feedback control. An IJP head, however, can remain several millimeters above the substrate, and hence is much less susceptible to destructive interactions with minor flaws in the object being fabricated.To embed printed transistors within fully-printed electromechanical devices, as mentioned previously, we must also develop a printable substrate which can support an OECT. Such a substrate material must be solution-processable, chemically compatible with all other materials used, electrically insulating, and must have a low surface roughness when printed atop a variety of other materials. We are in the process of evaluating several printable substrate materials.Materials Unmodified PEDOT:PSS is not a very good conductor, but many modifications used to increase its conductivity are beyond the scope of our system.

For example, ethylene glycol is often added to the PEDOT:PSS solution, improving the conductivity of the dried PEDOT:PSS film by a few orders of magnitude (Ouyang et al., 2004). It is difficult to process this solution in ambient conditions, though, because ethylene glycol does not dry in air, and heating the solution to hastendrying time releases toxic fumes (ethylene glycol is the active ingredient in antifreeze). Ultimately, we were unable to add ethylene glycol to our syringe deposition-printed OECTs, and their conductivity was significantly lower than that of previously reported OECTs (Nilsson et al., 2005). One advantage to IJP is that the very small volumes of liquid deposited by the IJP nozzle dry quickly, letting us integrate ethylene glycol into the PEDOT:PSS of future devices. We did add a small amount of DBSA (dodecylbenzene sulfonic acid) to the PEDOT:PSS solution (Lu et al., 2002) used to make both syringe deposition-printed and inkjet printed films.

We did add a small amount of DBSA (dodecylbenzene sulfonic acid) to the PEDOT:PSS solution (Lu et al., 2002) used to make both syringe deposition-printed and inkjet printed films. DBSA is a surfactant that helps the PEDOT:PSS form a more even film, and it caused us no noticeable problems.Annealing is also shown to benefit the conductivity of PEDOT:PSS (Kim et al., 2003), and heating the printed solution allows it to dry much more quickly, decreasing the total time required to print a single transistor. We decided to anneal the PEDOT:PSS film in air at a relatively low temperature (~150ºC) for half an hour, adding annealing functionality to our research platform by printing on top of a hot plate.We used an aqueous electrolyte solution for simplicity. Calcium, a divalent cation, had the potential to crosslink the PEDOT:PSS film and make it more robust (Ghosh et al., 1998), so we chose calcium chloride as our electrolyte salt.

Calcium, a divalent cation, had the potential to crosslink the PEDOT:PSS film and make it more robust (Ghosh et al., 1998), so we chose calcium chloride as our electrolyte salt. Preliminary experiments by our group showed no significant difference between the ionic conductivities of electrolyte solutions with different salts, so calcium chloride did not appear to be a better or worse choice than any other salt in this regard.Unfortunately, the silver ink used for the terminal contacts reacts electrochemically with the electrolyte solution. When any amount of the electrolyte made contact with the silver, a current spike appeared that was at least an order of magnitude higher than the currents we were trying to measure. Leaving a DC voltage on overnight across two silver ink electrodes bridged by the aqueous electrolyte solution revealed that a black film formed on one of the two electrodes.

Leaving a DC voltage on overnight across two silver ink electrodes bridged by the aqueous electrolyte solution revealed that a black film formed on one of the two electrodes. This film (perhaps silver chloride) indicated an electrochemical reaction between the electrolyte and the silver ink. We took care to ensure that the silver ink did not touch the electrolyte area of our OECTs, painting the contacts onto the transistor by hand for reliable placement, and we did not test transistors where the electrolyte leaked near or on the silver terminals. Our group is investigating a solution-processable carbon-based conductor that could alleviate these issues.One candidate for a suitable printable substrate is poly(4-vinylphenol) (PVPh) dissolved in 1- methyl-2-pyrrolidinone (NNMP). We would like our substrate to be compatible with IJP; one advantage to PVPh is that it was used previously as an insulating layer for an IJP-fabricated RC- filter circuit (Chen et al., 2003).

We would like our substrate to be compatible with IJP; one advantage to PVPh is that it was used previously as an insulating layer for an IJP-fabricated RC- filter circuit (Chen et al., 2003). Interestingly, PVPh was also used by Sandberg et al. in 2004 as the insulating layer of an OFET, with low operating voltages which they attributed to PVPh’s hygroscopic properties. However, these transistors had lateral features smaller than our setup can currently produce. Other printable substrate materials we are investigating include cyanoacrylate, epoxy, polyethylene glycol, and polycaprolactone.Experimental The geometry of the syringe deposition-printed transistors was designed using SolidWorks, a commercially available CAD program (see Figure 2). These exported files can be printed directly, with the only special software required being that which controls the printer.The active material, PEDOT:PSS (Baytron P, Bayer Corp.), was provided in an aqueous solution.

The solution was diluted in water (1 part Baytron P to 10 parts deionized water) to decrease the thickness of the dried film. Approximately 80µL/L of DBSA (dodecylbenzene sulfonic acid), a surfactant, was added to the solution to improve its wetting properties. The electrolyte used was an aqueous 1M CaCl2 (calcium chloride dihydrate, Mallinkrodt Chemical) solution. A material was needed to contain the drying PEDOT:PSS film and the electrolyte solution; we used GE Silicone II, a commercially available silicone adhesive designed for household applications.The transistors were printed using our group’s first generation multi-material freeform fabrication system, and all materials were extruded through syringes. More details on this research platform can be found elsewhere (Malone et al., 2004).

More details on this research platform can be found elsewhere (Malone et al., 2004). Annealing capability was added to decrease the drying time of the PEDOT:PSS film by printing while the glass substrate sat on top of a hot plate, which was turned on to heat the sample when necessary. Using these steps, we printed transistors:1) Print the thin silicone outline (Figure 2, shown on the CAD model in yellow) as a boundary for the PEDOT:PSS solution2) Print the PEDOT:PSS solution (shown on the CAD model in dark blue) within the silicone boundary as soon as the boundary dries3) Anneal the PEDOT:PSS at ~150ºC for 30 minutes 4) Print the thick silicone boundary (shown on the CAD model as transparent) to hold the aqueous electrolyte, and allow to dry5) Print the electrolyte solution (not shown) within the second silicone boundaryTransistors took a few hours to complete, with most of that time spent waiting for various components to dry.

Parallel fabrication was used to print four transistors on a single glass plate.Figure 2 – (Left) A 3D CAD model of a printable transistor. The yellow material is silicone (the second layer of silicone is partially transparent here to show underlying features), while the dark blue material is PEDOT:PSS. The liquid electrolyte (not shown) is held within the partially transparent boundary. (Right) A completed transistor set up for characterization, where only the electrolyte needs to be added (the grid size is 1cm2).The characterization equipment was in a different building than our freeform fabrication system, making it difficult to transport completed transistors for testing without the electrolyte spilling out of its shallow enclosure and making contact with a silver terminal, ruining the transistor for the purpose of our measurements. For ease of transport, we stopped printing these OECTs after Step 4, and added the electrolyte by hand at the testing site.

For ease of transport, we stopped printing these OECTs after Step 4, and added the electrolyte by hand at the testing site. However, we were able to successfully print an entire transistor complete with electrolyte, although we were unable to characterize it.The active area (the area immersed under electrolyte) of the source-drain channel of the final printed OECTs was approximately 3mm x 3mm, for a total channel area of 9 mm2, while the active area of the gate was approximately 6mm x 22mm, for a total gate area of 142 mm2. The ratio between the area of the gate and the area of the drain, then, is about 16:1, which is large enough to ensure that the gate is never over-oxidized.

The ratio between the area of the gate and the area of the drain, then, is about 16:1, which is large enough to ensure that the gate is never over-oxidized. A non-contact profilometry scan over a PEDOT:PSS film cast from undiluted PEDOT:PSS solution measured a film thickness of 5µm (the dilute films were too transparent to be characterized by this technique), so the PEDOT:PSS film cast from 1:10 diluted solution is estimated to be less than 1µm thick.In order to utilize IJP in future work, we have added an inkjet printer head to the Fab@Home Model 1. The IJP cartridge used is the HP 51604A, which uses a thermal activation mechanism for drop formation and expulsion. Information on controlling the firing mechanism for the printer and general information about ink jet printing can be found in a hobbyist kit from Parallax with a reference book (Gilliland, 2005).

Information on controlling the firing mechanism for the printer and general information about ink jet printing can be found in a hobbyist kit from Parallax with a reference book (Gilliland, 2005). The HP 51604A has 96 drops per inch resolution, and the average drop size from a given nozzle is 220pL (http://www.hp.com/oeminkjet/reports/TIJ1.0.pdf).Firing the nozzles of the IJP head is controlled by the Fab@Home’s microcontroller, which is connected to a non-retriggerable, monostable multivibrator circuit that generates short (~5µs) pulses. These pulses are amplified by a hex buffer, which in turn drives a Darlington transistor array, providing fast switching of the ink jet nozzle resistive heaters. A pulse of about 40 mJ of energy is delivered each time the nozzles fire, and this firing frequency can be varied using the Fab@Home software.

A pulse of about 40 mJ of energy is delivered each time the nozzles fire, and this firing frequency can be varied using the Fab@Home software. We can vary the density of droplets per unit area, as mentioned previously, to find the optimal firing frequency for each material printed.1 mmFigure 3 – IJP head on Fab@Home Model 1, printing PEDOT:PSS on a manually-applied epoxy substrate. (Inset) An optical micrograph of the printed PEDOT:PSS film on the epoxy substrate.For our preliminary experiments employing the IJP head on the Fab@Home Model 1 (see Figure 3), we deposited a solution with 2 mL PEDOT:PSS (either Aldrich, 2.8 wt% low conductivity grade or Aldrich, 1.3 wt% electronic grade) mixed with 9 mL water, .5 mL of DBSA, and .25 mL ethylene glycol (Mallinckrodt) onto various syringe-deposition printed substrates, as well as glass. Printed substrates included epoxy (3M, DP 100 Plus), cyanoacrylate (Loctite, super glue), and PVPh.

Printed substrates included epoxy (3M, DP 100 Plus), cyanoacrylate (Loctite, super glue), and PVPh. Specifically, the PVPh substrate was drop cast using a pipette onto a glass slide coated with a PET film, and allowed to dry, after which PEDOT:PSS was printed (with patterningperformed by hand) on top of the PVPh. PEDOT:PSS films ranged in size between 1 and 2 in2, with a thickness of approximately 1µm.Testing The syringe deposition-printed transistors needed additional modifications to connect to the characterization equipment. These modifications will be part of another printed component (like a wire) when the transistors are connected to a larger printed circuit. Conducting colloidal silver ink (Stan Rubenstein Assoc.) thinned 5:2 with toluene was painted onto the gate, source, and drain contacts (see Figure 2) and allowed to dry. Care was taken so that the silver remained far from the electrolyte solution.

Care was taken so that the silver remained far from the electrolyte solution. The silver ink provided a large contact area between the transistor and external probes, which were used to connect the transistor to the electrical measurement setup.The electrolyte was added to the transistors with an eyedropper at the testing site. It was allowed to sit for several minutes, giving its calcium ions a chance to cross-link the PEDOT:PSS film, before the transistors were tested. Transistors were characterized using two Keithley 2400 Source-Measure units interfaced with a LabVIEW VI. The source terminal of each OECT was grounded while voltages were applied to the gate and drain terminals.

The source terminal of each OECT was grounded while voltages were applied to the gate and drain terminals. The response of each transistor was determined by measuring the current flowing between the gate and source, referred to as the “gate current,” and between the drain and source, called the “drain current.” Resistivity data of PEDOT:PSS films were measured with a four-point probe, using a DC power supply and a multimeter to measure the voltage drop across a fixed distance of PEDOT:PSS film.Results and Discussion When a negative voltage is applied to the OECT drain terminal, while a positive voltage is applied to the gate, the transistor exhibits switching behavior, agreeing with previous reports (Nilsson et al., 2005). Figure 4 plots drain current versus drain voltage for a single transistor at various gate voltages.

Figure 4 plots drain current versus drain voltage for a single transistor at various gate voltages. There is a clear decrease in the drain current for increasing gate voltages.0.350.30.25)A u (t n e r r u C n i a r D0.20.150.1Vg = 0.2 VVg = 0.2 VVg = 0.1 VVg = 0.0 V0.0500.10.20.30.40.50.600.05Drain Voltage (V)Figure 4 – Drain current vs. drain voltage curves for a single printed transistor for various gate voltages. The transistor was measured twice at 0.2V applied to the gate, then once at 0.1V, and lastly once at 0V.With 0V applied to the gate, the transistor is in the “on” state. While on, the transistor in Figure 4 had approximately a 0.1MΩ channel resistance near 0V applied drain voltage, with similar on resistances noted for other OECTs we tested. Similar transistors from published literature are three orders of magnitude more conductive (Nilsson et al., 2005).

Similar transistors from published literature are three orders of magnitude more conductive (Nilsson et al., 2005). The main reason for this seems to be our inability to add ethylene glycol to the PEDOT:PSS and to anneal the film in an inert environment. However, with further optimization of film thickness and lateral dimensions, we could likely produce transistors that are more conductive.Applying a voltage of 0.2V to the gate reduces the drain current by about 50%, although the exact percentage depends on the current drain voltage. As the magnitude of the drain voltage increases, the drain current saturates, a behavior which occurs and is often exploited in FETs.

As the magnitude of the drain voltage increases, the drain current saturates, a behavior which occurs and is often exploited in FETs. (Why this occurs in OECTs as well is not obvious, but Bernards and Malliaras, forthcoming provides a good discussion of this effect.) It is reported (Nilsson et al., 2005) that the drain current of OECTs saturates for negative drain voltages but not for positive ones, which is why we used negative drain voltages for our preliminary tests (negative voltages applied to the gate reverse the direction of the electrochemical reaction, and quickly over-oxidize the channel).

Transistors that run off of negative drain voltages and positive gate voltages are unusual, and FETs are generally “off” with no applied gate voltage, rather than “on.” However, OECTs can still stand in for FETs in many cases, such as when they are used as active components in logic circuitry (Nilsson, 2005).The transistor was characterized twice at a gate voltage of 0.2V, then once at a gate voltage of 0.1V, and finally once at a gate voltage at 0V. The current degraded slightly when data were recorded twice in a row under the same conditions (the blue data points were taken before the pink data points), but this effect was negligible compared to the variation in current caused by 0.1V changes in gate voltage.

Since each data set took approximately 40 minutes to produce due to the time needed for the transistor to adjust to each voltage change (see Figure 5 for an example of the realtime data compiled to make Figure 4), it appears that these transistors are robust for at least hours and over several measurement cycles.Two of four fully-printed transistors on a single glass plate were functional, while the other two failed when their electrolyte solutions leaked and made contact with the silver terminals. The first transistor (the same from Figure 4) was about 50% less conductive than second transistor, but the second transistor, although still quite robust, also degraded somewhat more significantly than did the first. This indicates that the performance of these transistors is not necessarily consistent between different devices.

This indicates that the performance of these transistors is not necessarily consistent between different devices. More experiments should reveal the extent of this inconsistency, and devices made with our IJP head should also be more consistent with each other.Figure 5 shows simultaneously measured drain and gate current versus time curves for a single transistor. Both currents take at least a minute to settle when the voltage applied to the drain changes, and the same is true for a change in gate voltage. This switching speed seems very slow, but is acceptable for our first devices. In general, OECTs are slower than their FET counterparts; however, similar OECTs in the literature have been operable at 200Hz (Nilsson et al., 2002).

In general, OECTs are slower than their FET counterparts; however, similar OECTs in the literature have been operable at 200Hz (Nilsson et al., 2002). We should be able to optimize transistor speed by experimenting with the lateral dimensions and film thicknesses of the transistors, as large transistors with thick films likely switch more slowly than small and thin ones.The gate current spikes whenever there is a sudden voltage change, but within a minute or two it generally returns to zero. This indicates that the transistor works well, with no unwanted electrochemical reactions affecting the measured currents. When lower voltages are applied to the drain and gate, the gate (leakage) current is two orders of magnitude lower than the drain current. However, at higher voltages (≥0.5V), the gate current appears to saturate at a value significantly higher than zero.

However, at higher voltages (≥0.5V), the gate current appears to saturate at a value significantly higher than zero. It is difficult to operate these transistors with voltages any higher than 0.5V, and perhaps this is the threshold voltage for some sort of an electrochemical reaction between two of our components, or involves an impurity specific to our samples. Transistors described in the literature (Nilsson et al., 2005) did not report such difficulties.

time curves for the drain and gate currents of a printed transistor. Arrows indicate times when the drain voltage was changed to the voltage indicated. The gate voltage was held at 0.2V. (The blue line is the drain current, and the pink line the gate current.)We have favorable preliminary results from using our IJP head to print PEDOT:PSS films. The resistivity of the PEDOT:PSS film on PVPh was ~100 Ωm, while the film printed on glass was about an order of magnitude less resistive, with a resistivity ~10-1 Ωm. These figures are acceptable, as some loss of conductivity on a printed substrate is expected. Initial tests of PVPh suggest that it would make a good insulating substrate for our OECT devices.Conclusion Functioning transistors can be patterned using merely solution-processable materials that can be printed with our freeform fabrication system, as we have shown.

The significance of this accomplishment is due to the fact that this one system is also capable of producing batteries, actuators, structures, wiring, living tissue constructs, etc., and our work suggests that freeform fabricating printed transistors integrally with complex electromechanical devices is feasible. However, the characteristics, such as conductivity, switching speed, and reliability, of the transistors we have produced must be improved before we attempt to add other components to them. We plan to improve the performance of our transistors by optimizing their materialproperties, lateral dimensions, and film thicknesses. Our ink jet printer (IJP) head will give us more flexibility in patterning devices more reliably and at smaller scales than was possible with syringes, and may allow us to experiment with printing other types of transistors, like OFETs, in the future.Our current aqueous electrolyte may be easy to handle, but it has many disadvantages.

The low- viscosity solution can spill out of its enclosure or evaporate, unnecessarily lowering transistor yield and durability. Since PEDOT:PSS in solution is aqueous, damage to the PEDOT:PSS film is also a possibility when using an aqueous electrolyte. The dried film is not easily water-soluble, but a slight loss in conductivity occurred over about an hour of the PEDOT:PSS being immersed in electrolyte solution, and it is unlikely that our current transistors would have a particularly long shelf life (even ignoring electrolyte evaporation). It will be important to develop an electrolyte that has a high conductivity, is compatible with all our other materials, and does not evaporate.Ultimately, we wish to print working transistors on printable substrates. We are beginning to characterize how the properties of PEDOT:PSS are affected by a printable substrate, but will soon be ready to begin printing transistors on these substrates with the additional aid of our IJP head.

These ongoing improvements to our freeform fabrication system and to the performance of the transistors it prints, and work towards making printed transistors ready to integrate into larger printed systems, will greatly facilitate our efforts to create fully-functional electromechanical printed devices with a single machine.Acknowledgments This work was supported in part by the Cornell Center for Materials Research – Research Experience for Undergraduates program, which is supported by the National Science Foundation. R. H. acknowledges Sujit S. Datta for fruitful discussions.References Andersson, P., D. Nilsson, P. Svensson, M. Chen, A. Malmström, T. Remonen, T. Kugler, and M. Berggren. 2002. Active matrix displays based on all-organic electrochemical smart pixels printed on paper. Adv. Mater. 14:1460-64.Berggren, M., D. Nilsson, and N. D. Robinson. 2007. Organic materials for printed electronics.Nature Materials 6:3-5.Bernards, D. A., and G. G. Malliaras.

EEFORM FABRICATION OF STOCHASTIC AND ORDERED CELLULAR STRUCTURESJeffrey Lipton1, Mathew Boban1, Jonathan Hiller1, Hod Lipson1 1Cornell Computational Synthesis Lab, Department of Mechanical and Aerospace Engineering Cornell University, Ithaca NY, 14853Abstract Cellular materials provide a unique challenge to SFF technology. Such materials have unique properties of low mass, high strength, and good insulation properties. To produce such cellular structures, SFF systems require a designed microstructure with a feature size significantly lower than the resolution of the process. In this paper, we examine means of producing stochastic foams using the instability of a viscous thread and various methods for production of closed celled foams. These techniques allow for the production of foams without the need for pre-described cell structures.

These techniques allow for the production of foams without the need for pre-described cell structures. Such foams, when made from elastic materials can act as novel actuating materials.IntroductionCellular materials are ubiquitous in the modern world. Often they are used because of their low weight and outstanding insulating properties. They have been used as thermal, acoustic, and mechanical isolators. Varying the pore side and material density allows for the manipulation of stiffness, strength and other material properties. Cellular materials fall into two major categories: stochastic, and ordered. (1) Stochastic cellular materials include foams, sponges, and non-woven textiles. These produce an assortment of closed or open cells which do not follow any regular deterministic pattern. Non-woven textiles in particular are used in such varied applications as carpets, filters, surgical gowns, and textured surfaces.

Non-woven textiles in particular are used in such varied applications as carpets, filters, surgical gowns, and textured surfaces. Ordered cellular structures consist of geometric patterns such as truss structures, lattices, and honeycombs. (2) Traditional subtractive methods of manufacturing are generally incapable of directly producing such ordered microstructures.Background: Cellular StructuresSFF of cellular materials has focused on the production of ordered cellular structures. In order to produce stochastic foam, SFF technology required that the highly complex shape be completely specified in the geometry file (STL) prior to fabrication. Few CAD programs are capable of generating such complex shapes. David Rosen of Georgia Tech has developed tools for automatically generating regular mesostructures.

Few CAD programs are capable of generating such complex shapes. David Rosen of Georgia Tech has developed tools for automatically generating regular mesostructures. (3) However, production of such parts with a designed mesostructure requires the fabrication process to have a resolution much finer than the feature resolution of the mesostructure.Many applications of ordered cellular microstructure have used geometric processing techniques to create a lattice from an arbitrary geometry. These processes often rely on producing a strand with a width smaller than the width specified in parsing the geometries shape. This creates an open celled structure. In order to produce closed cellular structure, a fugitive material is required, or the shape must have each layer be geometrically identical. Fugitive material allows for each cell to be individually isolated. (4) A regular shape allows for theexterior boundary of the lattice to be sealed.

Fugitive material allows for each cell to be individually isolated. (4) A regular shape allows for theexterior boundary of the lattice to be sealed. Tissue scaffolds and piezoelectric sensors and actuators have been produced using these techniques. (5)Solid freeform fabrication of actuating materials has relied on electrically driven materials. Piezoelectric actuators which have been fabricated provide small directional displacement. (6). Ionomeric polymer-metal composite actuators have been successfully produced in the past and can be embedded in biocompatible geometries. However they have a directional actuation. (7) These constraints necessitated the development of a material with entrapped air. Such a material could expand and contract approximately isometrically when cycled through external pressures when made of elastic materials.Background: Viscous Thread InstabilityThe instability of viscous threads is familiar to anyone who has drizzled honey.

Thethreads buckle near the surface causing a coiling effect. This is known as the liquid rope coiling effect. According to Stephen Morris’s paper The Meandering Instability of a Viscous Thread, “The moving surface breaks one of the basic symmetries of the “pure” rope coiling problem, and has the effect of unfolding the coiling instability into a rich panoply of distinct bifurcations.” (8) These states can include: figure eights, meandering, translated coiling, and double coiling. The patterns are entirely predictable based on the nozzles width, relative speed, and height along with the material’s density, viscosity, and surface tension. Inducing these states in a robocasting system allow for the production of stochastic structures. The shapes and properties of the produced materials are deterministic based on the controlled parameters, but the microstructure produced is random.

The shapes and properties of the produced materials are deterministic based on the controlled parameters, but the microstructure produced is random. (9)Methods: ApparatusIn order to perform the experiments, a digital fabricator with adaptable path planning software was needed. A Model 2 Fab@Home Digital Fabricator with a syringe based deposition tool was used due to its ability to deposit a wide variety of materials, including elastic build materials and support materials. (2) FabStudio version 1 beta was used to generate the tool path information. It was selected for its versatility and the ease of modifications to path planning.(a)(b)Figure 2: Fab@Home Model 2 digital fabricator (a) and the vacuum chamber (b) used in experimentsIn order to demonstrate the ability of materials to actuate a single test geometry was selected. The primary test geometry was a rectangular prism of 20 mm by 20 mm by 10 mm.

The primary test geometry was a rectangular prism of 20 mm by 20 mm by 10 mm. This small geometry was selected for its ability to be quickly constructed and because volumetric actuation could be easily measured. A pressure chamber which can alternate between positive and negative 29 inches of mercury (98 kPa) from atmospheric pressure was used to actuate the closedcell foams. Using external pressure variations for the testing of the material allows without designing an interface to a pressure source.Methods: Foamed MaterialsInitial attempts focused on the creation of materials which would contain air when extruded. Such materials, when deposited, could turn any geometry directly into a foamed object. In order to entrap the air, liquid silicone was carbonated. Two pressure vessels were connected by a channel. In one vessel reactants of acetic acid and sodium bicarbonate were placed and allowed to react, producing carbon dioxide.

Two pressure vessels were connected by a channel. In one vessel reactants of acetic acid and sodium bicarbonate were placed and allowed to react, producing carbon dioxide. In the other vessel a large quantity of silicone was placed. Pressure was allowed to equalize between the chambers until the reactions in the reaction chamber had completed. The pressure was then released and the silicone was allowed to dry in bulk.Another method involved chemical reactions with the material. Sodium bicarbonate powder was rapidly mixed into liquid silicone. A solution of acetic acid and water was then mixed into the combined silicone/bicarbonate. The solution was mixed vigorously until the reaction had completed. The silicone was allowed to cure in bulk and a foamed structure was produced.Methods: Lattice Foamed MaterialCustom pathing algorithms which generate geometric distributions of air were attempted. These were based on an alternating regular rectangular lattice structure.

These were based on an alternating regular rectangular lattice structure. Initial attempts focused on sealing a region which has been sparsely filled. The primary pather decomposed the geometry into 3 regions: the bottom seal layers, the top seal layers, and the interior layers. The top and bottom regions were pathed using a conventional solid fill pather. The interior region used a custom pather which produced two concentric boundary paths to form a wall which sealed the layers form the exterior, and interior paths which left only one quarter of the space filled. The double wall ensured that a single error in the print process would not connect the interior air to the atmosphere. Dow Corning RTV Sealant 732, clear silicone was used as the build material since it is sufficiently elastic and prints reliably.This process relied on assumptions that can be made about the rectangular geometry of the prisms. The prism is 2.5-D; each layer had the same geometry as the previous layers.

The prism is 2.5-D; each layer had the same geometry as the previous layers. This ensured that the exterior solid boundaries were continuous and unbroken. The sparse filled region remained contained. In order to apply this method to a 3D geometry, the geometry would need to vary slowly, so that the outer boundary changes by no more than one path width inwards or outwards from layer to layer.Modifying this method allows it to be applied to an arbitrary geometry. If the geometry is sliced by three or more multiples of the materials native path height, it is possible to apply the sealed sparse planning algorithm to the slice. The top and bottom sub-layers are solid filled, and the intermediate sub-layers are sparse filled.

The top and bottom sub-layers are solid filled, and the intermediate sub-layers are sparse filled. (See figures 4 and 5)(b) Figure 4: An arbitrary geometry(a) can be sliced into sections which are integer multiples of the build materials native path height(b) which can then by pathed to ensure a closed sparse filled region(a)(a)(c)(a)(b)(c)(d)Figure 5: A Single slice of the geometry is divided into sub slices which are then pathed using either a sparse fill or solid fill pather. The solid fill(a)(d) sandwiches sparsely filled layers(b)(c)Method: Stochastic DepositionInstability of a viscous fluid flow can be induced by adjusting the flow and pathingcharacteristics. The values are manually tuned until the silicone enters the translated coiling state. The flow rate of the material is increased beyond its rate for construction of solid objects while the material’s deposition nozzle is elevated from the build platform beyond its normal solid construction height.

This allows the materials to spiral after the stream from the deposition head hits the build target. The relative height of the nozzle and the flow rate control the radius of the spiral of material. Path speed relative to fluid flow determines the linear spiral density. The density of the paths controls the spiraled material stream’s overlap between paths. (9) By using these modified flow values, it is possible to create an object where each layer consists of the translated coiled material. An object made from the coiling strand becomes a complex foam as each successful layer is added.Figure 6: typical regular spirals observed throughout much of the range for each parameterResults: Foamed MaterialsCarbonation of liquid silicone attempts resulted in several large (1mm diameter) bubbles in the surface of the silicone. The majority of the silicone remained unimpregnated with gas. Silicone was successfully foamed using sodium bicarbonate and additional acetic acid.

The majority of the silicone remained unimpregnated with gas. Silicone was successfully foamed using sodium bicarbonate and additional acetic acid. The uncured foam silicone could be extruded through a standard EFD syringe at 80 psi. The foamed silicone took significantly longer than normal to cure. The cured foam silicone was placed in the vacuum chamber and cycled through positive and negative 29 inches of Hg several times. The material expanded significantly during expansion stages, and contracted during positive pressure stages. However the material quickly lost its air content and the performance degraded too rapidly to be used as an actuator.Figure 9: A cross section of the foamed silicone created through sodium bicarbonate reactions shows it has a closed cellular structure.Results: Lattice Foam productionFailures of the conventional path planner and printing errors necessitated that the finalobject be tested for leaks.

The printed objects were submerged into water and pressure was applied to their surfaces. If air bubbles appeared, the entire object was coated in a thin layer of additional silicone and then allowed to dry. The process was repeated until no bubbles could be observed. The prisms on average required two coatings of silicon after printing. The path planner used generated paths which did not completely fill the top and bottom solid fill layers. These gaps were due to an error in the algorithm which expanded the layers boundaries to create the walls. The test prisms produced using this method placed in the vacuum chamber expanded to when exposed to -29 inches of mercury from atmosphere.

The test prisms produced using this method placed in the vacuum chamber expanded to when exposed to -29 inches of mercury from atmosphere. Prisms when exposed to +29 inches of mercury from atmosphere.Figure 8: the Path planner deposits material (grey) to build the object but the errors in path planning leave unintended gaps (white) which need to be filled by a bulk coating post processing.Results: Stochastic DepositionTranslated coiling of the viscous thread produced open celled foams. Changes in the spiral width of the translated coils produced foams of different pore sizes and surface roughness. An arbitrary geometry (space shuttle) was printed using two different spiral widths. The geometry was processed using the modified values of path width and height of the translated coils.

An arbitrary geometry (space shuttle) was printed using two different spiral widths. The geometry was processed using the modified values of path width and height of the translated coils. The sizes of the pores were of the same order as the resolution of the printing process.(a)(b)(c)(d)Figure 9: A material which can be solid printed (a) can be used to create a stochastic foam (b) by using viscous thread instability. By adjusting the properties of the viscous flow, a single geometry could be made into foams of various properties (c) (d).Discussion:Previous research in SFF allowed for the direct fabrication of designed ordered cellular structures. Complex cellular features and structures required the pores to be orders of magnitude larger than the process resolution and complex CAD programs to generate the intended geometry (3).

Complex cellular features and structures required the pores to be orders of magnitude larger than the process resolution and complex CAD programs to generate the intended geometry (3). Simple lattice cellular structures allowed for pores of the same order as the print process; they could not be used to generate closed cellular structures of arbitrary geometry. The processes used to create the lattice structures required a fugitive support material, or relied on the regularity of the shape. (4)The geometric processing algorithm used to generate tool paths for FDM machines described earlier shows how a single material can be used to generate shapes with entrapped air. The algorithm overcomes traditional geometric limitations by using a process of generating sub layers of identical boundaries with contain the cellular structure while allowing for the overall shape to vary.

Such material distributions can be used to create bulk volumetric actuators.Volumetric actuators are a novel form of Solid Freeform fabricated actuators. They are the first isometrically expanding additively manufacturable actuator. While the direct deposition of a closed cell foamed elastic material would be the ideal means of creating an external pressure driven volumetric actuator, they are not suitable for the creation of internally driven actuators. Additionally the closed celled materials made from chemical reactions were not as robust as the geometrically planned materials. It should be possible to use the geometric processing methods to create pneumatic actuators that are powered by internal pressure changes.The deposition of material by means of viscous thread instability allows for the creation of materials of stochastic cellular structure.

The pores are inherently of print process order and do not require any additional computation to translate a solid geometry into a stochastic foam of the desired shape. This demonstrates the ability to create a foamed object without the need for descriptions in CAD or complex geometric processing.The ability to freeform fabricate foams of arbitrary shape but controlled pore size and void fraction could allow for novel application of SFF. Catalytic converters, filters, and any other application of either non-woven threads or foams could not benefit from the novel abilities of SFF. Additionally it is possible to use a single material to create gradations in mechanical properties by varying the print parameters used. (9) Textures could be applied to additively manufactured parts without the need to have the texture features be orders of magnitude larger than the process resolution.Conclusion:The techniques described here allow for the fabrication of various cellular structures.

A single material can be used to generate closed and open celled regular lattices. It can be used to produce stochastic open celled foams using viscous thread instability. Or it can be directly foamed to create closed cell stochastic foams. These processes enable new applications for solid freeform fabrication without the need for CAD tools or complex algorithms. The closed celled lattice has already demonstrated the ability to create a novel volumetric actuator.Acknowledgements:This work was supported in part by a National Science Foundation Graduate ResearchFellowship and NSF Creative-IT grant 0757478Works Cited 1. Investigation of Solid Freeform Fabrication Processes for the manufacture of Parts with Designed Mesostructure.

e Freeform Fabrication of Structurally Optimized and Complexly-Shaped Metal Tubular ComponentsDr. Kevin Chaite Rotheroe, Freeform Design & Manufacturing Research Studio, New York School of Architecture & Building Research Council, University of Illinois, Urbana-ChampaignAbstractThe service conditions of many structural frames composed of tubular metal components would ideally warrant the use of high strength-to-weight ratio components with shapes and internal geometries that respond to context-specific structural requirements. Commercially available and emerging solid freeform fabrication technologies can be utilized to indirectly or directly manufacture metal tubular structural components with optimizing features that cannot otherwise be manufactured. The results of prototyping experiments demonstrating the viability and potential of this application of additive manufacturing will be presented.

The results of prototyping experiments demonstrating the viability and potential of this application of additive manufacturing will be presented. This presentation will discuss successful prototype 356 aluminum and 316 stainless steel internally reinforced freeform tubular components manufactured indirectly using expendable patterns made by selective laser sintering and 3D printing. The application of laser and metal powder based freeform fabrication technologies that provide superior material properties will also be discussed, especially in terms of requirements for multi-axis deposition and sophisticated path planning software, and the implications of voxel- or layer-based functionally gradient materials.1.0 IntroductionThe term tube is conventionally understood to refer to a hollow lineal component with a constant cross-sectional shape and size. Constant cross-sectional configuration, however, is not inherent in the definition.

Constant cross-sectional configuration, however, is not inherent in the definition. Complex hydroformed automotive chassis components, for example, are referred to as being tubular, and are created by deforming conventional metal tubes.Conventional roll-formed or extruded metal tubes can be considered to exhibit generally optimized strength-to-weight ratios achieved by a combination of hollowed simple cross- sectional geometry and particular material properties.1.1 Classifying Structural OptimizationThe author has defined two general classes of structural optimization as it pertains to components utilized in structural frames. Externally optimized components are those that generally maximize strength-to-weight or strength-to-material usage ratios with features that are visible on the exterior of the component. Rolled steel sections produced to industry standards such as angles or ‘I’ beams are common examples exhibiting simple cross-sectional geometry.

Rolled steel sections produced to industry standards such as angles or ‘I’ beams are common examples exhibiting simple cross-sectional geometry. Non- hollow extrusions also fall into this category. Examples of much more complex externally optimized components include subtractively formed bulkheads such as those found in aircraft fuselage frames: such components can feature a wide variety of complex product-specific web, flange, and rib configurations. A few research and/or commercial organizations are presently47developing hybrid additive and subtractive methods for making complex externally optimized components such as bulkheads. AeroMet Corporation, for example, has developed gantry- configured 3-axis deposition of titanium onto webs that are pre- (or post-) cut from titanium sheets.

AeroMet Corporation, for example, has developed gantry- configured 3-axis deposition of titanium onto webs that are pre- (or post-) cut from titanium sheets. This deposition process forms ribs and flanges, and required surface accuracy is achieved by CNC machining the deposited material.Internally optimized components are those that generally maximize strength-to-weight or strength-to-material usage ratios using features that are formed internally and/or are not visible on the exterior of the component, except possibly at open ends or via external shape, as with a hydroformed member. Conventional tubes are very basic internally optimized components in that their hollowness is not visible unless holes or perforations have been created.

Conventional tubes are very basic internally optimized components in that their hollowness is not visible unless holes or perforations have been created. Extrusion is frequently used to manufacture much more complex internally optimized components, and many of these components, such as those found in aluminum building fenestration systems or in the chasses of manufacturing equipment, can feature highly complicated cross-sectional geometry. The extent, however, to which complex internally optimized components such as extrusions can obtain maximum strength-to-weight ratios and respond to context-specific service conditions, is limited by formal constraints inherent in available manufacturing methods.

Additive processes, especially direct metal deposition methods, have significant potential to circumvent these constraints and enable further optimization of metal structural frame components.The general constraints on product form inherent in established manufacturing processes are listed below. A taxonomy of all available manufacturing processes for making metal structural components is presented in Figure 1. This taxonomy conventionally classifies manufacturing processes as mass-conserving or mass-reducing, but then sub-classifies processes based on the geometric constraints they impose on product design.1. Equipment-based constraints. The formal output (e.g. possible product size) of a given manufacturing process is limited by at least one piece of available required equipment.2. Motion-based constraints:(cid:131) Rotational Motion results in axially-constrained product geometry.

Motion-based constraints:(cid:131) Rotational Motion results in axially-constrained product geometry. (cid:131) Lineal Motion (centralized applied force) results in products with a constant cross-sectionalconfiguration (e.g. extrusion, rolling).(cid:131) Surface Distributed Lineal Motion (distributed applied force) can only produce partswithout undercut or re-entrant features (e.g. stamping and forging).(cid:131) Multi-Axis Motion cannot access or create significant internal features when utilized inconjunction with subtractive manufacturing methods.3. Tooling-based constraints. The need to remove mandrels, re-usable patterns, dies, re-usable molds, and required fixtures, as well as the components (products) formed with these tools, significantly limits the potential geometry of products, or, in other words, the extent to which optimizing features can be incorporated into designs.4. Process Based Constraints. Workpiece access and material removal requirements (e.g.

Axially-Constrained Formation or Deformation Processes are those that create products with axial or rotational motion. Processes that extrude or roll material through dies or molds generate products that are both axially constrained and constant in cross section. Subractive turning processes create axially-constrained products, but these may vary significantly in cross section. 2. Surface-Constrained Formation and Deformation processes are those that form material within the constraints of part-specific surfaces. Indirect Tool-Dependent Processes form parts within the external surfaces of the component or the internal surfaces of its tooling. In subtractive formation, material is removed up to the external surfaces. Surface-Constrained processes are also limited (constrained) by the need for surface-oriented fixturing or jigs (subractive processes) or die and mold parting surfaces. 3.

Surface-Constrained processes are also limited (constrained) by the need for surface-oriented fixturing or jigs (subractive processes) or die and mold parting surfaces. 3. Fundamental manufacturing process for making industry standard structural metal materials and stock products (e.g. sheets, plates, rolled sections, tubes). 3a. A process by which standard tubular products are made. 3b. Centrifugal Casting is used to make standard tubular products, but it is more often used to make manufacturer-specific products. 4. Mass Deformation refers to the deformation of a preformed and component-specific ingot containing a pre-determined quantity of metal in a pre-engineered configuration. Hot Rolling of plates and sheets is, like the production of rolled sections, often directly linked to a continuous casting operation. 5. Metal 3D Printing is being developed by the ProMetal Divison of ExtrudeHone Corporation, licensee of MIT’s 3D Printing patent.

5. Metal 3D Printing is being developed by the ProMetal Divison of ExtrudeHone Corporation, licensee of MIT’s 3D Printing patent. tooling system for injection molding and other tools requiring complex internal passages that cannot be made by conventional methods (e.g. for conformal cooling). ProMetal is developing larger build volume machines that will make the production of structural components technically viable for some applications, although it will likely be many years before the process is commercially viable for architectural structures. 6. Research into the use of metal sheets in Laminated Object Manufacturing is ongoing at Helisys, Inc. and The Warwick Manufacturing Group (UK). Metal LOM is not presently commercially viable and should be considered a process in R&D. The most significant technical obstacles are removal (de-cubing) of metal from cavities and the surface dimensional accuracy resulting from aliasing (stair-stepping).

The most significant technical obstacles are removal (de-cubing) of metal from cavities and the surface dimensional accuracy resulting from aliasing (stair-stepping). While research into variable cutting angles per or within a given layer (metal sheet) is ongoing, resolution of this issue will still result in products with a degree of surface abstraction which will necessitate finish machining operations for many applications. 7. The Replicast Process, invented simultaneously at The Castings Development Centre (UK) and Waukesha Foundry (WI), combines the dimensional and surface accuracy of Investment Casting with the size capacity of bonded sand molding processes. A relatively thin ceramic shell mold is positioned in a flask which is then filled with a sand/binder mixture, vibrated and compacted, and placed under a vacuum during metal pouring. 8. The V-Process was invented in Japan for making architectural products such as gratings and ornamental panels.

8. The V-Process was invented in Japan for making architectural products such as gratings and ornamental panels. It is now largely used for making mechanically and or structurally functional components as well as housings. 9. Squeeze Casting, which is also known as Semi-Solid Forging, is a hybrid process combining the metallurgical advantages of forging with the greater design freedom permitted by casting. Molten metal solidifies under pressure in metal molds and the casting is then ejected in a manner similar to injection molding. 10. In Spray Casting, which is also known as the Osprey Process, molten metal is sprayed onto a rotating mandrel as it passes through a controlled chamber. The process is used to manufacture seamless tubes and pipes. 11. Hydroforming operations can be either axially-constrained or surface constrained. It is commercially viable as a rapid1.

The process is used to manufacture seamless tubes and pipes. 11. Hydroforming operations can be either axially-constrained or surface constrained. It is commercially viable as a rapid1. Axially-Constrained Formation or Deformation Processes are those that create products with axial or rotational motion. Processes that extrude or roll material through dies or molds generate products that are both axially constrained and constant in cross section. Subractive turning processes create axially-constrained products, but these may vary significantly in cross section. 2. Surface-Constrained Formation and Deformation processes are those that form material within the constraints of part-specific surfaces. Indirect Tool-Dependent Processes form parts within the external surfaces of the component or the internal surfaces of its tooling. In subtractive formation, material is removed up to the external surfaces.

In subtractive formation, material is removed up to the external surfaces. Surface-Constrained processes are also limited (constrained) by the need for surface-oriented fixturing or jigs (subractive processes) or die and mold parting surfaces. 3. Fundamental manufacturing process for making industry standard structural metal materials and stock products (e.g. sheets, plates, rolled sections, tubes). 3a. A process by which standard tubular products are made. 3b. Centrifugal Casting is used to make standard tubular products, but it is more often used to make manufacturer-specific products. 4. Mass Deformation refers to the deformation of a preformed and component-specific ingot containing a pre-determined quantity of metal in a pre-engineered configuration. Hot Rolling of plates and sheets is, like the production of rolled sections, often directly linked to a continuous casting operation. 5.

Hot Rolling of plates and sheets is, like the production of rolled sections, often directly linked to a continuous casting operation. 5. Metal 3D Printing is being developed by the ProMetal Divison of ExtrudeHone Corporation, licensee of MIT’s 3D Printing patent. tooling system for injection molding and other tools requiring complex internal passages that cannot be made by conventional methods (e.g. for conformal cooling). ProMetal is developing larger build volume machines that will make the production of structural components technically viable for some applications, although it will likely be many years before the process is commercially viable for architectural structures. 6. Research into the use of metal sheets in Laminated Object Manufacturing is ongoing at Helisys, Inc. and The Warwick Manufacturing Group (UK). Metal LOM is not presently commercially viable and should be considered a process in R&D.

and The Warwick Manufacturing Group (UK). Metal LOM is not presently commercially viable and should be considered a process in R&D. The most significant technical obstacles are removal (de-cubing) of metal from cavities and the surface dimensional accuracy resulting from aliasing (stair-stepping). While research into variable cutting angles per or within a given layer (metal sheet) is ongoing, resolution of this issue will still result in products with a degree of surface abstraction which will necessitate finish machining operations for many applications. 7. The Replicast Process, invented simultaneously at The Castings Development Centre (UK) and Waukesha Foundry (WI), combines the dimensional and surface accuracy of Investment Casting with the size capacity of bonded sand molding processes. A relatively thin ceramic shell mold is positioned in a flask which is then filled with a sand/binder mixture, vibrated and compacted, and placed under a vacuum during metal pouring. 8.

A relatively thin ceramic shell mold is positioned in a flask which is then filled with a sand/binder mixture, vibrated and compacted, and placed under a vacuum during metal pouring. 8. The V-Process was invented in Japan for making architectural products such as gratings and ornamental panels. It is now largely used for making mechanically and or structurally functional components as well as housings. 9. Squeeze Casting, which is also known as Semi-Solid Forging, is a hybrid process combining the metallurgical advantages of forging with the greater design freedom permitted by casting. Molten metal solidifies under pressure in metal molds and the casting is then ejected in a manner similar to injection molding. 10. In Spray Casting, which is also known as the Osprey Process, molten metal is sprayed onto a rotating mandrel as it passes through a controlled chamber. The process is used to manufacture seamless tubes and pipes. 11.

The process is used to manufacture seamless tubes and pipes. 11. Hydroforming operations can be either axially-constrained or surface constrained.

Plasma-Arc and Laser Cutting are commonly used for producing job-specific components such as custom plates or web members for metal fabrications. 2. Continuous Mechanical 3D Cutting is generally only used for making small components and connectors in tensile structures or fenestration support structures, and, of these, axially-constrained formation is much more common. More complex versions of such small structural parts are sometimes investment cast. 3. Structural Investment Castings are generally only used in architecture for small connectors (e.g. glazing bosses) or brackets in job-specific secondary structural systems such as those supporting a window wall. Manufacturers of specialty glazing systems sometimes provide system-standard investment castings. 4. While job-specific structural Sand Castings (mostly iron) were used extensively in the latter half of the 19th century for components in compression (e.g.

4. While job-specific structural Sand Castings (mostly iron) were used extensively in the latter half of the 19th century for components in compression (e.g. columns), their use rapidly declined with the advent of stronger standard rolled sections. The use of sand castings for both primary and secondary structural components in buildings , while still not common, has increased significantly in the past approximately fifteen years, especially in the United Kingdom. 5. Centrifugal Casting was used more frequently in architectural structures in the 19th Century. It is sometimes used as an alternate means of producing stock or job-specific tubular (typically column) components. Centrifugal castings are much denser than statically-poured castings and are conseqently stronger (but not as strong as rolled tubes). 6. Continuous Casting is a basic ingot-producing process. Continuous Castings are typically fed directly into hot rolling operations to produce standard optimized sections. 7.

6. Continuous Casting is a basic ingot-producing process. Continuous Castings are typically fed directly into hot rolling operations to produce standard optimized sections. 7. Fundamental process for making standardized steel products for the construction industry. 8. Extrusion is not typically used for making components for primary structural systems in buildings because it cannot compete economically with hot rolling for steel products (higher temperatures increase die cost and wear). It is, however, the dominant technology for making aluminum structural frames for fenestration assemblies. 9. Mandrel Swaging of tubes is occassionally used for making tapering ends for larger tubular steel structural components. 10. The two-dimensionally-constrained bending of standard optimized structural sections for job-specific steel frame components is fairly common in architecture. 11.

10. The two-dimensionally-constrained bending of standard optimized structural sections for job-specific steel frame components is fairly common in architecture. 11. Three-dimensional bending of plate products is quite rare in architecture, although commonplace in the marine industry. 12. Sheet products are generally accepted to be those less than 6mm (1/4”) thick and plate products to be 6mm or greater in thickness. While planar stamping of sheet products is a common means of manufacturing architectural cladding components, in other industries sheetmetal is frequently stamped into complexly-shaped structural frame components as well as protective enclosures, as is the case with unitized automotive structures. 501. Plasma-Arc and Laser Cutting are commonly used for producing job-specific components such as custom plates or web members for metal fabrications. 2.

501. Plasma-Arc and Laser Cutting are commonly used for producing job-specific components such as custom plates or web members for metal fabrications. 2. Continuous Mechanical 3D Cutting is generally only used for making small components and connectors in tensile structures or fenestration support structures, and, of these, axially-constrained formation is much more common. More complex versions of such small structural parts are sometimes investment cast. 3. Structural Investment Castings are generally only used in architecture for small connectors (e.g. glazing bosses) or brackets in job-specific secondary structural systems such as those supporting a window wall. Manufacturers of specialty glazing systems sometimes provide system-standard investment castings. 4. While job-specific structural Sand Castings (mostly iron) were used extensively in the latter half of the 19th century for components in compression (e.g.

10. The two-dimensionally-constrained bending of standard optimized structural sections for job-specific steel frame components is fairly common in architecture. 11. Three-dimensional bending of plate products is quite rare in architecture, although commonplace in the marine industry. 12. Sheet products are generally accepted to be those less than 6mm (1/4”) thick and plate products to be 6mm or greater in thickness. While planar stamping of sheet products is a common means of manufacturing architectural cladding components, in other industries sheetmetal is frequently stamped into complexly-shaped structural frame components as well as protective enclosures, as is the case with unitized automotive structures. 5042.0 Current Design ContextAt present, the majority of structural metal frame components are designed, specified, or selected from standard configurations to accommodate areas of maximum stress or loading.

The other portions of a given member are therefore inherently over engineered and/or sub-optimally designed. For example, the cross sectional size and configuration of an aluminum extrusion used as a beam is determined by the portions of its length that will transfer the most significant structural loads. Other portions of the same component may experience much smaller loads or transfer different types of forces. These portions would ideally have a different configuration, but the component’s overall geometry is dictated by more critical loads incurred elsewhere.While this relationship between design and manufacturing is to varying degrees universally true, it is especially true in architectural and civil applications where most structural frames are one- off configurations composed of industry-standard component members.

Unlike other industries where production volumes can justify product-specific tooling costs, architects, civil and structural engineers frequently have to design custom context-specific structural frame solutions using components whose tooling and production costs are distributed via standardization across multiple related industries. Figure 2 is an analytical taxonomy of manufacturing processes for making structural metal components, with those methods presently used in architectural and civil applications highlighted.

Figure 2 is an analytical taxonomy of manufacturing processes for making structural metal components, with those methods presently used in architectural and civil applications highlighted. Mass customization enabled by large scale indirect and direct metal solid freeform fabrication has the potential to enable the circumvention of design constraints that are presently universal in the building design and construction industry.3.0 Classifying Internal Structural OptimizationFor the purposes of this research it was necessary to assess all extant processes for manufacturing metal structural components in terms of their capacity to produce structurally optimized and complexly-shaped tubular components. It also became necessary to qualify and classify the extent to which some of these processes possess such a capacity. The results of this assessments and classifications are presented in Figure 4.

It also became necessary to qualify and classify the extent to which some of these processes possess such a capacity. The results of this assessments and classifications are presented in Figure 4. The varying extent to which those manufacturing processes with a capacity to produce internally-reinforcing features could produce such features resulted in distinct categories of internally-reinforced free-form tubes:1. Freeform tubes with no internal reinforcing. Either additive or subtractive methods can be used to make expendable patterns or reusable tooling for the indirect manufacture of free- form tubes with variable or constant wall thickness. Expendable patterns are used in conjunction with various expendable ceramic mold casting processes. (Examples are shown in Figure 7b.) Hard tooling is utilized, for example, to hydroform complexly-shaped tubes.2. Freeform tubes with limited internal reinforcing.

(Examples are shown in Figure 7b.) Hard tooling is utilized, for example, to hydroform complexly-shaped tubes.2. Freeform tubes with limited internal reinforcing. These are freeform tubes with internal reinforcements that do not bridge between opposing tube walls, or, in other words, internal surface reinforcements that can be formed within the constraints of reusable pattern removal requirements (e.g. as with bonded sand mold casting). Such surface reinforcement or ribbing is common on non-tubular components and housings, and could be introduced on the inner (or outer) surfaces of expendable or reusable patterns for freeform tubes.3. Freeform tubes with extensive internal reinforcing.

Freeform tubes with extensive internal reinforcing. This class of freeform tubes highlights the type of optimization that it was hypothesized could be achieved using additive51technology to directly (utilizing laser (or electron beam) and metal powder based deposition processes) manufacture such a component, or to indirectly manufacture it using an expendable pattern casting process.

The specific features that can be incorporated into the design of a complexly-shaped and internally reinforced tube are:a) Complex corrugated tube walls.b) Internal bracing that bridges between opposing external furrows or surfaces, in somecases via a core that follows the tube’s curvilinear trajectory.c) Curvilinear openings within the internal braces that demonstrate the possibility ofimproved strength to weight ratios in these features.d) Relatively thin walls that demonstrate the possibility of improved overall strength to weight ratios with the use of internally-reinforcing features enabled by additive formation.4.0 Indirect Manufacture of Structurally Optimized and Complexly-Shaped Metal Tubular Components with Solid Freeform FabricationA variety of experimental prototypical freeform tubes with extensive internal reinforcing were designed in ProEngineer with the features enumerated in item 3 above.

All of these features, as well as connective features and the overall composition of the tubes were parametrically variable so that these computer models could be readily modified in response to various structural and manufacturing simulation analyses. The indirect manufacture of these prototypical shape optimized internally reinforced tubes was then tested: pattern components were created using 3D printing (Z Corporation Z402 machine) and selective laser sintering (both Castform™ and Trueform™ materials). These self-registering pattern components were then glued together and transported to aerospace foundries where gating and other pattern features were added. These prototypes were successfully investment cast in both 356 aluminum and 316 stainless steel. They received various heat treatments in order to maximize their strength.

These prototypes were successfully investment cast in both 356 aluminum and 316 stainless steel. They received various heat treatments in order to maximize their strength. Two of these prototypes are 1820 mm in curvilinear length, but most are approximately 600 mm in curvilinear length, such as the one shown below in Figure 3.Fig. 3a & 3b. Portion of an A356 structurally optimized and complexly-shaped tube with extensive internal reinforcing features.

Electro-Chemical Machining (ECM) is used to machine forging dies from high-strength alloys and to form complex cavities in aerospace components. ECM technically has the capacity to produce complex solid and tubular forms, but these would be limited in size by the work range of the machine set-up (typically for small parts) and expensive: Part-specific tooling requires significant production volume or peformance criteria for the process to be economically viable . 2. Multi-Axis Mechanical CNC Milling can easily produce solid complexly-shaped metal parts, but complex tubes of any significant depth (beyond 2 times the depth of the bit - and chuck depending on opening size) cannot be produced unless they are longitudinal “halves” for making complex tubes as welded or assembled fabrications. 3.

3. Standard tubular products can be hydroformed into either axially-constrained or complexly-shaped variable cross section tubes, the later being increasingly used in the making of automobile chasis components. Equipment and tooling costs are high. 4. The use of Swaging to make variable section tubular components is inherently constrained by the need to remove the required mandrel after formation. 5. Swaging operations can utilze either all rotating dies (inherently axially constrained) or the combination of a stationary and a rotating die, which permit the formation of non-axial complexly-shaped solid parts, although this is less common. 6. All Ceramic Mold casting processes can produce complexly-shaped tubular components with internal (and external) features if necessary shell bridges and access holes are accommodated, and especially if patterns are generated by Solid Freeform Fabrication technologies.

However, Poured Ceramic and Plaster Mold processes are generally used only for small components, and designs for use with Patternless Investment Casting must accommodate the build volumes of SLS machines or assembly from SLS mold components. The size capacity of Investment Casting has increased significantly in recent years and parts up to 60” or 80” are now successfully cast. 7. Bonded Sand Molds, Dry Sand Molds and the Replicast Process can be used to make complexly-shaped tubes with internal (and external) features, but these processes are significantly limited relative to Ceramic Mold Casting processes for two reasons: a). Pattern removal during mold making will not accommodate undercuts, and; b). Patterns are typically made using multiple-axis CNC machining/routering which would require subdividing the pattern into many components (discreet machining operations) to accommodate significant internal complexity.

Even then it would be difficult or impossible (depending on the design) to match the capability of Solid Freeform Fabrication. 8. Process can inherently make complexly-shaped tubes with internal features, but the design must accommodate the removal of unsintered or unbound powder, and products are very limited in size by the build volumes of various machines. 9. Research into the use of metal sheets for LOM is ongoing at Helisys, Inc. and The Warwick Manufacturing Group (UK). The process could make some complexly-shaped tubular designs, but its potential is very limited by the need to remove (de-cube) internal material. 10. Permanent Mold Casting can produce complexly-shaped tubular parts if sand cores are used. This practice is fairly common. Fig. 4.

10. Permanent Mold Casting can produce complexly-shaped tubular parts if sand cores are used. This practice is fairly common. Fig. 4. Analytical Taxonomy: Manufacturing Complexly-Shaped Metal Components537775.0 Architectural Applications of Complexly-Shaped and Structurally Optimized Metal Tubular ComponentsComplexly-shaped and structurally optimized metal tubular components are being designed by the author’s architectural practice, Freeform Design & Manufacturing Research Studio, and manufactured for use in full size architectural structures. The components shown in Figures 6 and 7, for example, are prototypes for the conservatory structure shown in Figure 5. This structure was designed as an assembly of discreet parametrically-variable components in ProEngineer, and these components were in turn subdivided as required into self-registering pattern components that will be manufactured both subtractively (CNC milled from high density foam) and additively using 3D printing.

The pattern components will then be assembled and bonded together with foundry glue prior to being manufactured using a variety of hybrid sand and ceramic mold casting processes. The components were digitally analyzed both structurally and in terms of castability (i.e. with MAGMA), and modified as required (i.e. per stress analysis, metal flow, and shrinkage). The model shown in Figure 5 was also designed as a kit of parts in ProEngineer, and these parts were made with stereolithography, urethane castings made from RTV tools and stereolithography patterns, and laminated object manufacture.Fig. 5. Structural Model of a Conservatory Fig. 6. Component PatternsFig. 7a. A356 Prototype 7b.

5. Structural Model of a Conservatory Fig. 6. Component PatternsFig. 7a. A356 Prototype 7b. 316 Stainless prototypes, as cast and powder coated.546.0 Notes on Biomimetic Structural FormThe notion of optimizing the strength-to-weight ratio and/or strength-to-material usage ratio of a tubular structural component by optimizing its geometry (design) per anticipated product- specific service conditions can be considered a biomimetic approach to structural design enabled by solid freeform fabrication. This is because such an approach attempts to replicate some of the material efficiencies found in biologically generated structures. The author has defined two fundamental classes or approaches to biomimetic design:1.

The author has defined two fundamental classes or approaches to biomimetic design:1. The direct imitation of biologically generated structural efficiencies can be considered a scientific approach that includes: (cid:131) replicating structural composition, which is always relative to other biological functionsand naturally occurring service conditions;(cid:131) imitating material efficiency by minimizing and functionally varying or distributingmaterial usage(cid:131) generally maximizing overall strength to weight ratio per natural precedent.2. Aesthetically-driven or interpretive organic structural form, which is in part or whole an artistic approach to biomimetic design.Design for commercial products (as opposed to performance specified, entirely non-aesthetic products) involves the reconciliation of the scientific and artistic approaches to interpreting natural structures.

The balance can vary greatly, especially in architecture, and indeed this is reflected in the conservatory design shown here. However, the research introduced in this paper (e.g. per Figures 3a and 3b above) is focused solely on maximizing engineered “biomimetic” structural efficencies, not only for architectural purposes, but for a wide variety of small and large scale applications.6.1 Cautionary Notes on Biomimetic Structural Form1. Nature is generally highly efficient (i.e. economical). But, there is a distinction between natural economy and human economy. After satisfaction of service conditions (including embedded safety factors), economic efficiency (cost) governs most commercial design (e.g. the retail warehouse is a high strength-to-weight ratio construct maximizing enclosed volume whilst generally minimizing material usage – while a more efficient structure might be possible, the cost of manufacturing the required components does not add sufficient value).2.

The higher the degree of material efficiency relative to function, the greater the extent to which a structure might be considered “biomimetic,” regardless of its form.3. The distinction between natural economy and human economy exists because nature has more sophisticated manufacturing (material forming) processes. SFF (especially in conjunction with FGM) and the emerging capacity to manufacture biologically (i.e. grow components and materials) are major steps in the ever-increasing human capacity to replicate nature. The commercialization of this increasing capacity is (as it always has been) constrained by cost benefit realities: the value added by new means of achieving conventional or previously impossible solutions must be sufficient to justify increased cost.

557.0 Concluding DiscussionFree-form tubular components with extensive internal reinforcing have been successfully indirectly manufactured using a variety of expendable mold casting methods and expendable patterns made with 3DP & SLS. These prototypes range up to 1820 millimeters in curvilinear length. In addition, a wide variety of internal and external “biomimetic” geometric configurations have been developed. The quantification (measurement) of strength-to-weight ratio improvements possible with these geometries is ongoing, and benchmark geometric configurations are being developed for experimental prototyping and comparative analysis using a variety of direct metal processes. Powder based direct metal SFF processes that utilize surrounding powder for support (SLS & 3DP/ProMetal) can presently best accommodate complex biomimetic tubular designs.

Powder based direct metal SFF processes that utilize surrounding powder for support (SLS & 3DP/ProMetal) can presently best accommodate complex biomimetic tubular designs. The only major constraints on design with these processes are the need to remove powder from internal cavities and the limited (but improving) achievable tensile strengths.Current research is focused on small scale applications, and assumes the capacity to manufacture complex internal features while also matching or exceeding the properties of wrought alloys.

This assumption recognizes the fact that many of the possible benefits of structurally optimized and complexly-shaped metal tubular components cannot be fully realized utilizing casting processes, even though expendable patterns with the required geometries can be produced using SFF: the limitations of achievable material properties, especially tensile strength, even with appropriate heat treatments, can result in failure to fully capture the benefits of improved geometry, especially for service conditions requiring tensile performance.Various laser or electron beam and metal powder based deposition methods, including LENS™ and other proprietary processes, can presently best satisfy the material strength requirements of the designs. These processes, however, must address other significant issues and limitations in order for the manufacture of structurally optimized and complexly shaped tubular components to become truly viable. These issues include:1.

These issues include:1. Circumvention of the need for support structures in inaccessible areas via controlled multi- axis deposition.2. Most path planning software for direct metal deposition processes cannot presently accommodate complex geometric configurations. This is a very significant hurdle.3. Accuracy and surface quality must be improved. Surface roughness can lead to fatigue cracking and catastrophic structural failure. Some service conditions will require CNC milling of surfaces to eliminate this potential. Internal surfaces are therefore problematic and suggest the need for hybrid metal deposition and subtractive processes. The path planning requirements for such conditions are a significant hurdle.4.

The path planning requirements for such conditions are a significant hurdle.4. Capturing the full benefits of biomimetic designs for freeform metal tubes is ultimately dependent on realization of significant capacity to achieve functionally-graded materials – controlled local deposition of alloys via at least layer-based FGMs, but preferably voxel- based material specification and voxel-cluster material delivery.1 The term Laser Engineered Net Shaping is used in Figures 1,2 and 4 to generically describe laser or electron beam and metal powder deposition processes.56.

eeform Fabrication of Ti3SiC2 StructuresD.J. Dcosta and W. Sun(1) Department of Mechanical Engineering and MechanicsT. EI-Raghy Department of Materials EngineeringDrexel University Philadelphia, PA 19104This paper introduces a three-stage process to fabricate highly dense Ti3SiC2 structures. The properties of Ti3SiC2 material, the synthesis of the ceramic powder, the procedure involved in the 3-stage fabrication process and the preliminary results on fabricating fully dense Ti3SiC2 structures are presented. The characterization and microstructure evaluation of the mechanical, morphological and structural properties covering the compressive strength, Vickers micro hardness, damage tolerance, thermal shock, shrinkage and porosity of Ti3SiC2 structures printed using the 3-stage process are presented.1.

IntroductionQuite a few researches have used freeform fabrication techniques to directly prototype functional ceramic parts in order to obtain more freedom in the design and utilization of ceramic structures [1]. Studies include photopolymer resin curing by ultraviolet light employing SLA and solid ground curing [2], thermal fusion of powders by laser scanning used in the SLS process [3], adhesion of powders by disbursal of binder through an ink-jet used in three-dimensional printing [4-6], compacting and sintering of powders as demonstrated in the freeform powder molding technique [7], extrusion of thermoplastics through a fixed diameter nozzle in the FDM process [8], using an extension of the 3D Printing technique to build objects of variable density by controlling microstructure and composition [9], and Multiphase Jet Solidification (MJS) [10].

Another technique involved a further development to the technology of Inkjet Printing, wherein, the binder in the printing cartridge was replaced by a ceramic suspension, pressurized in a continuous ink jet printer [11,12]. Similar techniques such as Robocasting [13], involving a computer-aided deposition of highly concentrated ceramic colloidal slurries directly forming complex shapes without the use of molds or dies have also been researched upon.

Additionally, other similar techniques relying on the basic inkjet printing technology are the overprinting technique [14] using inkjet printing of electro ceramics such as lead zirconate titanate (using a piezoelectric actuator to eject the ink); electro photographic powder deposition [15] technique employing the deposition of charged powder particles layer-by- layer, resulting in the precise deposition of particles to the desired shape; electromagnetic ink jet printing [16] employing the ejection of ceramic inks via an electromagnetic valve printing head and dispersion of ceramic inks onto a substrate using an ultrasonic vibrating horn have also been experimented upon.(1) Corresponding author, Tel: 215-895-5810; Fax: 215-895-2094; Email: sunwei@drexel.edu203The objective of this paper is to present our preliminary studies on using Layered Printing, compacting, and sintering processes to fabricate fully dense functional Titanium Silicon Carbide (Ti3SiC 2) structures with the comparable properties to the Ti3SiC 2 bulk reported in [17- 20].

The contents of this paper are presented as follows. Section 2 briefly reviews the properties of Ti3SiC 2 and the ceramic powder preparation. Section 3 describes the 3-stage process. The fabricated samples are presented in Section 4. The results of structural characterization, such as the compressive strength, Vickers micro hardness, damage tolerance, thermal shock, shrinkage and porosity and microstructure evaluation, of Ti3SiC2 structures are presented in Section 5. Summary, conclusion and discussions are presented in Section 6.2. Ti3SiC2 - Properties and Powder PreparationTitanium Silicon Carbide, a layered ternary ceramic compound was processed in bulk, by Dr. M. W. Barsoum and Dr. T. El- Raghy at Drexel University. Polycrystalline bulk samples of Ti3SiC 2 were synthesized by reactively hot pressing Titanium, Graphite (Carbon) and Silicon Carbide powders at a pressure of 40 MPa and temperature of 1600 oC for 4 hours at a temperature rate of 10 oC/min. [18 - 20].

[18 - 20]. The bulk sample obtained as a result of the solid phase reaction was milled to obtain a ceramic powder. As a part of the powder preparation process pure Ti3SiC 2 powder is thoroughly mixed with a water-soluble binder and dried to form clinker. The resulting clinker is crushed in a powder mill. As a result of this entire process fine Ti3SiC 2 powder particles coated with a film of binder are obtained. Further, the powder was sifted through a –325 mesh sieve to obtain an average grain size of less than 40 m m. This batch of powder was used as the build powder in our study.Ti3SiC 2 possesses a unique combination of mechanical, electrical and thermal properties of both metals and ceramics. Like metals it is an excellent electrical and thermal conductor, easily machinable, relatively soft, good thermal shock resistance and behaves plastically at higher temperatures, and like ceramics it is oxidation resistant and refractory in nature[18,19].

It has an electrical conductivity of 4.5 x 106 ohm-1 m-1 that is twice that of pure Titanium (2.3 x 106 ohm-1 m-1). The thermal expansion coefficient, in the temperature range of 25 to 1000 oC, the room temperature thermal conductivity and heat capacity are 9.2 x 10-6 oC-1, 43 W/m K and 588 J/kg K respectively, these being similar to that of Titanium namely, 31 W/m K and 523 J/Kg K. Its density of 4.5 gm/cc is roughly half the density but almost double the stiffness of Nitrogen based super-alloys. Ti3SiC 2 is relatively soft with Vickers Hardness of 2-5 GPa but elastically stiff and easily machinable.A remarkable property of Ti3SiC 2 carbide is its easy machinablility in spite of its high stiffness. The layered structure of material and its ductility contributes to direct machinability to close tolerances without any post machining sintering, whereas in other cases the ceramic needs to be sintered after machining resulting in an additional shrinkage.

Ti3SiC 2 is the only carbide that can be cut manually using a hacksaw. Due to its good electrical properties it is also easily shaped and cut using Electric Discharge Machining. In addition, Ti3SiC 2 retains its mechanical properties at high temperatures and is stable at temperatures of 1800 oC with excellent high temperature mechanical properties [20]. There is no effect on the bend strength of Ti3SiC 2 when heated to 1400 oC and quenched. This ceramic actually increases in strength and exhibits excellent resistance to thermal shock, whereas other ceramics have a drastic reduction in strength204when subjected to the same conditions of quenching. The structural properties of samples developed using the 3-stage integrated process have been evaluated in Section 5.3. The Three-Stage Fabrication ProcessThe fabrication of Ti3SiC 2 structure is divided into three primary stages [21]: 1) solid freeform fabrication, 2) pressure compacting, and 3) Sintering.

The Three-Stage Fabrication ProcessThe fabrication of Ti3SiC 2 structure is divided into three primary stages [21]: 1) solid freeform fabrication, 2) pressure compacting, and 3) Sintering. Flow chart of the 3-stage process is described in Figure 1. In the first stage, the part is built using the layered Printing technique employing the powder deposition and liquid binding technique to print the part as a layered structure, wherein the Ti3SiC 2 powder is used as the build material. In this process the STL file generated from the corresponding CAD model is inputted to the prototyping system to generate the process path. The green part is built by deposition of Ti3SiC 2 powder and binder spray (as per the cross-section at the particular slice) layer by layer till the entire part is built. At this stage the green part has been freshly developed and has a low structural strength.

At this stage the green part has been freshly developed and has a low structural strength. The second stage is to compact the green part by applying a high pressure on the green part to increase the density and the strength of the green part. The last stage of the process is to sinter the compacting part in an electric tube furnace at high temperature in a flowing Argon atmosphere (inert atmospheric conditions are specified to prevent any oxidation).Figure 1: Three-stage fabrication process for Ti3SiC 2 structures4. Fabrication of Ti3SiC2 SamplesTo understand the effect of the compacting pressure on the structural properties of the Ti3SiC 2 samples, two sets of samples were processed. One set of samples was fabricated using the integrated 3-stage process, i.e., SFF, Compacting, and Sintering. While the other set was sintered directly without Compacting after SFFP.

One set of samples was fabricated using the integrated 3-stage process, i.e., SFF, Compacting, and Sintering. While the other set was sintered directly without Compacting after SFFP. The samples without the compacting process205were fabricated with layer thickness of 0.0035”, 0.007” and 0.01” and the binder concentration of 7 – 40%. Standard procedures were used to prepare samples for structural and optical characterization.Figure 2 shows different Ti3SiC 2 samples fabricated through the 3-stage process, such as a 6-tooth sprocket, turbine wheel and hollow tube (Figure 2), and the master sample for test and characterization (Figure 3). Figure 2a to 2c show the samples at different process stages, namely, SFF green parts (2a), compacted parts (2b), and sintered parts (2c). The master sample shown in Figure 3 was used as the test specimen. Test samples for mechanical property evaluation were derived from the master sample.

The master sample shown in Figure 3 was used as the test specimen. Test samples for mechanical property evaluation were derived from the master sample. The layer thickness for fabricating the master sample was fixed at 0.0035”. The densities achieved for samples fabricated using the 3-stage technique were over 99%.Figure 2: Parts fabricated using the 3-stage processFigure 3: Master sample (layered structure) fabricated using the 3-stage process2065. Mechanical Testing and Microscopic EvaluationThe mechanical testing and microscopic evaluation of the prototyped Ti3SiC 2 samples were conducted to evaluate the compressive strength, Vickers micro hardness, damage tolerance, thermal shock, shrinkage and porosity (Figure 4). A summary of the results is presented in this section. Comparisons have been made with the available data reported in the references.

A summary of the results is presented in this section. Comparisons have been made with the available data reported in the references. It needs to be pointed out that the Ti3SiC 2 powder batch used in our samples contained about 30 – 40% more TiC (Titanium Carbide) than the powder used in reference [20], and hence the results reported here are higher than the reported data due to the presence of the excessive amount TiC.Figure 4: Mechanical characterization and microscopic evaluation of Ti3SiC 2 test samplesThe test specimens were cut from the master sample (Figure 3) using an EDM. Parallel machining was ensured to minimize any irregularities in dimensions of the test samples. For the evaluation of micro hardness and optical microstructure, samples cut from the master sample.

Parallel machining was ensured to minimize any irregularities in dimensions of the test samples. For the evaluation of micro hardness and optical microstructure, samples cut from the master sample. were subjected to a grinding sequence of 300, 500 and 1200 grit Silicon Carbide paper and later polished to a mirror finish on 3 m m and 1 m m diamond suspension using a Struers Rotopol automatic polishing machine.The specimens for compression testing were cut from the master sample in the X and Z axes of the printing development, and for damage tolerance and thermal shock in the X and Y axes of the printing development. The sides of the specimen were ground on 800 and 1200 grit Silicon Carbide paper to remove burrs and flashes. This was done to minimize the possibility of error during the four point bend test.

The sides of the specimen were ground on 800 and 1200 grit Silicon Carbide paper to remove burrs and flashes. This was done to minimize the possibility of error during the four point bend test. The specimens for the damage tolerance tests were marked with an indentation using a Vickers Indenter, on the tensile side of the specimen with indentation loads of 100, 200 and 300 N. It was ensured that the indentation207was in the region of maximum bending of the specimen. The samples for the thermal shock tests were heated upto pre-quench temperatures of 400 oC, 800 oC and 1200 oC.5.1 Compression StrengthThe compression tests were conducted on test specimens measuring 3mm x 3mm x 6mm cut along the X and Z axes of the printing using a servo-hydraulic testing machine (MTS System, Eden Priarie, MN) at a constant crosshead speed of 0.01 mm/s.

Friction and mutual sliding between the Testing Machine platens and the specimen surfaces was minimized by the application of anti- friction grease. The breaking stress in compression was calculated as:s = Load / AreaThe Load – Displacement curve is presented in Figure 5. It is observed that the deformation in compression of the test specimen is linear to failure and does not exhibit any plastic deformation behavior, with the failure being that of a brittle fracture. The results of the compression tests were calculated based on an average of five measurements. The average breaking strength in compression at a breaking load of 10.3 KN was found to be 1144.44 MPa in the XY plane of the printing, and at a breaking load of 11.03 KN was 1225 MPa in the Z plane, both at a strain rate of 0.0016/s. The observed difference in strength along the two independent planes or axes leads us to believe that the strength property is anisotropic due to the Layered Printing process.

The observed difference in strength along the two independent planes or axes leads us to believe that the strength property is anisotropic due to the Layered Printing process. Although the micrographs (presented later) do not reveal any texture in the specimen, we believe that there is a process- induced anisotropy in the printed structure. The reported compressive strength of compacted and sintered (without SFF) coarse-grained Ti3SiC 2 samples at room temperature is 720 MPa [20]. The test results conform to the reported compressive strength values for compacted (without SFF) and sintered coarse-grained Ti3SiC 2 samples, with a proportional increase observed due to the presence of 30 – 40% of TiC.Figure 5: Load - Displacement curve and compressive strength2085.2 Vickers Micro HardnessThe micro-hardness test was conducted on polished sample surfaces using a Leco M-400 (St. Joseph, MI) diamond indenter at loads of 1, 3, 5 and 10 N.

Joseph, MI) diamond indenter at loads of 1, 3, 5 and 10 N. The diagonal values of the indentations were measured, an average of the diagonal lengths (Figure 6) was calculated and a conversion was obtained from the corresponding Vickers charts to derive the Vickers Micro Hardness.D1D2MicrographSchematicFigure 6: Indentation obtained using Leco M400 Micro Hardness TesterThe micro hardness values have been presented as an average of five measurements. The hardness obtained for the samples was in the range of 12.3 – 9.4 GPa for an indentation load of 1 – 10 N (Figure 7). Further, the hardness seemed to have stabilized at 9 GPa, i.e, Vickers Hardness at increased loads will result in the stabilized value.

Further, the hardness seemed to have stabilized at 9 GPa, i.e, Vickers Hardness at increased loads will result in the stabilized value. The Vickers hardness of compacted and sintered (without SFF) Ti3SiC 2 samples has been reported to be in the range of 8 - 4 GPa for samples obtained from fine grained Ti3SiC 2 with load increasing from 1 N to 100 N, and for coarse-grained samples the reported range is 5 - 4 GPa for the same range of indentation load [20]. The hardness value asymptotically stabilizes at 4 GPa for loads greater than 100 N. This corresponds to the similar trend observed in the SFF samples. The results obtained conform to the established values for Ti3SiC 2.

This corresponds to the similar trend observed in the SFF samples. The results obtained conform to the established values for Ti3SiC 2 . An increase in VHN, as we believed, is due to the presence of excessive 30 – 40% of TiC imparting greater hardness to the samples.Figure 7: Vickers Micro Hardness for Ti3SiC 2 samples2095.3 Damage ToleranceFor the damage tolerance tests the specimens were prepared as per ASTM C1161 type A (25mm x 1.5 mm x 2 mm) along the X and Y directions of SFF Layered Printing. Indentations on these specimens were performed using a Vickers indenter under loads of 100, 200 and 300 N, with one diagonal of the indenter parallel to the length of the specimen. The flexural strength after indentation was determined by a four point bend test as per the procedure in ASTM C 1161 and ASTM C 1211 using a four point flexural fixture. The cross head speed was set for 0.01 mm/s, this corresponds to an average strain rate of 0.0066/s.

The cross head speed was set for 0.01 mm/s, this corresponds to an average strain rate of 0.0066/s. To remove any surface defects that could act as a crack initiation site under load, the tensile surfaces of the specimens were polished with 1200 grit Silicon Carbide prior to the indentation.The results for damage tolerance are reported as an average of five measurements taken for each indentation load. The results reported here are for indentation loads of 100, 200 and 300 N. Damage tolerance was recorded as a measure of the breaking strength in a four-point bend test. The retained flexural strength as a result of indentation decreases from 387.33 MPa at no load to 251.67 MPa at 300 N for the sample length along the X-axis of the printing dierction, and decreases from 385 MPa to 242.92 MPa for the sample length along the Y-axis, as shown in Figure 8.

The flexural strength of compacted and sintered (without SFF) Ti3SiC 2 samples has been reported to be in the range of 300 MPa at no load to 240 MPa for a 300 N indentation load in the case of coarse-grained samples [20]. The flexural strength decreased steadily with the increase in the indentation load. This trend is reflected in SFF samples, and the results conform well to the reported data for coarse-grained Ti3SiC 2.Figure 8: Damage Tolerance for Ti3SiC 2 samples5.4 Thermal ShockFor the thermal shock resistance, specimens conforming to ASTM C1161 type A (25mm x 1.5 mm x 2 mm) along the X and Y directions of the printing were placed in a furnace at the pre-quench temperatures of 400, 800 and 1200 oC, under flowing argon (to prevent the formation of any oxide layers) for 10 minutes. The samples were then immediately quenched into an210ambient temperature water bath.

The samples were then immediately quenched into an210ambient temperature water bath. The quenched samples were tested according to the ASTM C1161 and ASTM C 1211 using a four point flexural fixture. All four point bend tests were conducted using a servo- hydraulic MTS Systems testing machine running at a cross head speed of 0.01 mm/s, corresponding to a nominal strain rate of 0.0066 /s. In both cases, thermal shock resistance and damage tolerance, the flexural strength was calculated using the following formula:s = 3 P (l1-l2) / (2 B W 2 )are the outer and the inner span, respectively, B is the where P is the load at fracture, l1 and l2 specimen width (2 mm), and W is the specimen thickness (1.5 mm). The four-point bend fixture used had an outer span of 20 mm and inner span of 10 mm. The fixture rollers were free to rotate, but constricted laterally in this setup.

The four-point bend fixture used had an outer span of 20 mm and inner span of 10 mm. The fixture rollers were free to rotate, but constricted laterally in this setup. The mutual friction between the sample and the rollers was eliminated by the application of a very thin film of anti- friction oil on the roller interface.The results for thermal shock are reported as an average of five measurements taken for each temperature condition. Thermal shock has been reported as a measure of the flexural strength of the sample under the four-point bend test. The tests were conducted at room temperature, 400 oC, 800 oC and 1200 oC. The retained flexural strengths for the sample length along the X-axis of the printing are in the range of 387.33 – 211.67 MPa and 395 – 174.58 MPa for the sample length along the Y-axis of the printing. The test samples thermally shock up to 400 oC and the Module of Rapture further stabilizes and is retained up to 1200 oC.

The test samples thermally shock up to 400 oC and the Module of Rapture further stabilizes and is retained up to 1200 oC. Beyond 800 oC, an increase in post quench strength is observed in both sample sets, as shown in Figure 9. The flexural strength of compacted and sintered (without SFF) Ti3SiC 2 samples subjected to thermal shock are about 260 MPa for a temperature of 1400 oC for coarse grained samples and ranging from 560 – 170 MPa over a temperature range of 500 oC to 1100 oC for fine grained samples [20]. It is observed that coarse grained samples retain flexural strength up to a temperature of 1400 oC and are thus thermal shock resistant, whereas fine grained samples thermally shock between 750 oC and 1000 oC [20].Figure 9: Thermal Shock Resistance for Ti3SiC 2 samples211The test results reflect the reported results of thermal shock resistance and the increasing trend of post quench strength at temperatures of 800 oC for coarse-grained Ti3SiC 2.

Additionally, it is also observed that the flexural strength along the X-axis of the printing is greater than that along the Y-axis. This is mainly due to the layered nature of the Layered Printing technique. The published results of the Module of Rapture for damage tolerance and thermal shock for Ti3SiC 2 samples are as per tests conducted without regard for the axis of development. The presented values as per this research demonstrate the anisotropy in strength arrived at as a result of the Layered Printing process.5.5 Porosity and ShrinkageThe physical shrinkage of each of the samples was measured and calculated as a function of the linear dimensions D, of the individual parts. The shrinkage was calculated as a percentage after each process, namely, Layered Printing, Compacting, and Sintering.Shrinkage = (D pre-process – D post-process ) / D pre -processThe percentage porosity is an average of 5 independent representative regions of the sample surface under test.

The sample that was derived form the master sample (subjected to the post RP processes of compacting and sintering) had a density of greater than 99%. For the samples developed by varying the binder concentration and layer thickness during the Layered Printing process and directly sintering without compacting, the porosity obtained from the surface calculations of the micrographs show an average porosity of 50 – 60%. This shows that compacting is the dominant process in improving the density of the part. This is quite noticeable comparing the micrographs between Figures 10 and Figure 12. The optical microscope was used to measure the cross-sectional porosities of each sample. The supporting software provides image capture, threshold and density slice tools by which percentage of black and white phases are measured directly.

The supporting software provides image capture, threshold and density slice tools by which percentage of black and white phases are measured directly. The average porosity for each sample is calculated as a mean value for the data retrieved by the random choice of at least 5 representative points on each sample surface.The shrinkages after each stage of the integrated process were calculated as a percentage for the sprocket, turbine wheel and tube as listed in Table 1.

The average shrinkage measured after the compacting was in the range of 15% and after sintering was about 30%.Table 1: Shrinkage data for parts developed using the integrated 3-stage processSprocket / Turbine: Geometric FeaturesLength Width HeightPre- Compact inches 1.52 1.52 0.50Post- CompactCompactinches % change 1.33 1.33 0.4212.50 12.50 16.00Post SinteringCompact & Sintering inches % change 1.06 1.06 0.3630.07 30.07 27.30Tube: Geometric FeaturesLength OD IDPre- Compact inches 1.00 0.62 0.30Post- CompactCompactinches % change 0.81 0.49 0.3019.00 20.97 0.00Post SinteringCompact & Sintering inches % change 0.69 0.42 0.2230.70 31.61 28.332125.6 Optical MicroscopyMicrographs of the various samples representing dense and porous structures were acquired using the Olympus PM-3G Optical Microscope.

From the micrographs shown in Figures 10, we observe that the part density is in excess of 99% in the planes XY, XZ and YZ through the 3-satge process. The sample structures exhibiting greater than 99% density show a homogeneous distribution of grains without any texture within the powder matrix as shown in Figures 11. The main cause of high density and structural homogeneity is due to the compacting process. In the case of the samples exhibiting porous structure it is observed that the average porosity measured is about 50 - 60% as shown in Figure 12. From the micrographs we observe that the sintering without compacting leads to binder burnout and results in the part being porous. This explains the voids (dark regions) on the micrographs as shown in Figures 12.

This explains the voids (dark regions) on the micrographs as shown in Figures 12. Further, the increased binder concentration increases the possibility of sintering of a lower quality.Figure 10: Micrographs of XY, XZ and YZ planes of the master sample With density greater than 99% (200 m m)Figure 11: Micrographs of the cross-section of the master sample with density greater than 99% at X 500 and X 1000 magnificationFigure 12: Micrograph of sample with porosity at 60% (porous structure) (200 m m)2136. Summary and ConclusionsThis paper introduces an integrated process of Layered Printing, Compacting, and Sintering to fabricate highly dense Ti3SiC2 structures. Experiments were conducted to study the mechanical characteristics and the effect of the processing parameters on the morphological and structural properties of the fabricated Ti3SiC2 structures. The results of the mechanical characterization and the microstructure evaluation of Ti3SiC2 structures have been presented.

The results of the mechanical characterization and the microstructure evaluation of Ti3SiC2 structures have been presented. The resulting sintered parts were achieved with a dens ity of greater than 99%. As a post processing surface finishing the parts can be subjected to a grinding and polishing sequence to obtain a mirror finish. As a demonstration of the machinability of the SFF parts, EDM was used to cut the test samples and polished them to obtain required 1m m surface finish. This integrated process was used to freeform fabricate structures with complex shapes, as shown in Figures 2. A density of greater than 99% was achieved for each of these parts.The Load – Displacement curves show a linear to failure deformation behavior of the Ti3SiC 2 specimens.

A density of greater than 99% was achieved for each of these parts.The Load – Displacement curves show a linear to failure deformation behavior of the Ti3SiC 2 specimens. The results of the mechanical characterization for the Ti3SiC 2 samples also reveal an anisotropy in the mechanical properties, for example, the compressive strength, damage tolerance and thermal shock properties are greater along a certain axis of the printing development. Although the SEM micrographs show a homogeneous grain distribution of Ti3SiC 2 powder without any texture within the part matrix, we believe that the layered printing process contributes to the anisotropy in the prototyped structure.

The results from the tests for compressive strength, hardness, damage tolerance and thermal shock are comparable to the reported data for Ti3SiC 2, considering the effect of the presence of excessive 30 – 40% TiC in the Ti3SiC 2 powder used in the fabrication of samples under test.The samples fabricated using the 3-stage process achieved a density of greater than 99%, while samples without the compacting had a porosity in the range of 50 – 60%. This shows that compacting is the dominant process leading to the high structural density. The average shrinkage measured after compacting was in the range of 15% and after sintering was 30%. The main cause of high density and structural homogeneity is due to the compacting process. A certain amount of warpage was observed after the printing. We believe that this is mainly due to the capillary forces distributed in the powder matrix during the printing process.

A certain amount of warpage was observed after the printing. We believe that this is mainly due to the capillary forces distributed in the powder matrix during the printing process. The maximum warpage occurred along the Y-axis (longest dimension) of the printing and was further enhanced after the compacting process.In comparison to other Solid Freeform Fabrication methods for prototyping dense and functional structure, the developed 3-stage process is simplicity and cost-effective. However, the process of a larger scale and complicated geometry needs to be developed further to possibly eliminate the compacting process, thus to avoid the part collapse and to retain the designed geometry. The process-induced anisotropy and its relation with the processing parameter and part geometry need also to be further investigated.References1.J. G. Heinrich: “New developments in the solid freeform fabrication of ceramic components”, Ceramic Forum International, Vol. 76, No. 5, 1999, pp.

EEFORM FABRICATION OF ZINC AIR BATTERIES WITH TAILORED GEOMETRY AND PERFORMANCEMegan E. Berry, Evan Malone, Hod Lipson Mechanical and Aerospace Engineering Cornell University Ithaca NY, USAReviewed, accepted August 26, 2005AbstractThis work is focused on the freeform fabrication of complete zinc-air batteries. This method of production gives great freedom in the geometry and construction of the battery, allowing tailoring of the output characteristics, and the possibility of embedding a customized battery within a larger integrated freeform fabricated device. Our batteries utilize a gelling agent to prevent phase separation in the zinc anode and catalyst layers, permitting the use of nozzles down to 1.3mm in diameter. Polyvinyl alcohol is utilized for the separator layer, which replaces the un- printable paper separator used in commercial batteries.

Polyvinyl alcohol is utilized for the separator layer, which replaces the un- printable paper separator used in commercial batteries. With various freeform batteries we have achieved a specific capacity of 60mAh/g of zinc, an average power 7.25mW, and continuous service life of 63h, all with a load of 100 ohms. The specific capacity of our freeform batteries is about 1 order of magnitude lower than that of commercial zinc-air batteries, although under different test conditions. We have investigated the effect of cell active surface area on performance for a cylindrical cell-geometry, and have produced a flexible, two-cell battery with unusual geometry. The tailoring of performance and geometry possible with freeform fabrication will be of great value in the design of optimized smart devices with unusual geometry, portable electronics, and prototypes.IntroductionBatteries are a necessary component in all electronic systems.

By manufacturing batteries using freeform fabrication we open the door to creating completely freeform fabricated circuits as well as batteries integrated within other systems in a way not possible with current manufacturing techniques. Interest in flexible batteries is rising for embedded applications in smart cards and RFID (Akashi et. al, 2002). For example, PowerPaper Ltd. (2005) is manufacturing thin, flexible, alkaline cells through a lamination process. In addition, research in the area of direct-write electronics (Piqué and Chrisey, eds., 2002) has led to write-able electrochemical cells; however, these are embedded into integrated circuits using traditional, non-freeform, IC manufacturing. In prior work, we presented the first freeform fabrication of a complete, functional zinc-air battery (Malone et. al., 2004). Our past and current work remains the only example of the freeform fabrication of complete batteries.

al., 2004). Our past and current work remains the only example of the freeform fabrication of complete batteries. In the following, we present improvements to the materials and processes which have allowed us to freeform fabricate batteries in several geometries, and with tailored functionality. We developed new formulations for the zinc anode, catalyst, electrolyte, and separator materials which provide a satisfactory level of cell performance, while being easily deposited. We improved the manufacturing planning software of our system in order to accelerate and improve the quality of fabrication of multiple material devices such as batteries. We produced and tested a flexible, two-cell zinc-air battery with an unusual geometry, which demonstrates the design freedom which is possible using our freeform fabrication system. We have collected data on the effect of battery geometry on performance.

We have collected data on the effect of battery geometry on performance. To isolate the effect of geometry, we fabricated cylindrical batteries with differing surface areas and a standard mass of active material. The performance data from these will allow us to extend the predictive performance model of Mao and White (1992). A mature predictive295model and freeform fabrication capability will allow the size, shape, and performance of a cell to become free design parameters which can be customized for the specific geometry and functional requirements of a product. This will permit us to design and fabricate customized batteries for use in prototypes, portable electronics, and other novel applications. Battery design can be adapted to fit the existing product, rather than having the limited characteristics of commercial batteries dictate product design.

Battery design can be adapted to fit the existing product, rather than having the limited characteristics of commercial batteries dictate product design. An explanation of zinc-air chemistry will be followed by a presentation of the material formulations we have developed, the process of freeform fabrication for a battery, our testing procedure, and finally the batteries we have produced and their performance data.Background A cell is defined as a device that converts the chemical potential energy between its anode and cathode materials into electrical energy by means of redox reactions: reduction (electron gain) at the cathode, oxidation (electron loss) at the anode. A battery comprises one or more connected cells. The components of a zinc air cell are the anode terminal, cathode terminal, zinc anode, separator, catalyst, and electrolyte (Figure 3). The anode and cathode terminals are charge conductors which collect and transfer charge to external loads.

The anode and cathode terminals are charge conductors which collect and transfer charge to external loads. The zinc anode and the oxygen cathode are the two key reactants in the cell. The separator provides electrical insulation between the anode and cathode to prevent internal shorting, but must be permeable to electrolyte to allow ionic current. The catalyst accelerates the rate of reaction in the cell, improving power output. The electrolyte, in this case aqueous potassium hydroxide, KOH, provides a medium for ionic transport within the cell, and a source of ions. The basic chemical reactions are:At Air Cathode: ½ O2 + H2O + 2e →2OH- At Zn Anode: Zn → Zn2+ + 2eZn2+ + 2OH- → Zn(OH)2 Zn(OH)2 → ZnO + H2O______________________________________ Overall:Zn + ½O2 → ZnO : E0=1.65 VFormulation of MaterialsZinc anode pasteIn creating the zinc anode paste we were required to balance several concerns.

The zinc particles needed to be fully suspended in the electrolyte to provide a reliable flow rate without clogging or phase separation. We wanted to reduce hydrogen gas formation in the anode paste which causes problems with gas pockets in the loaded syringes and a loss of power output in a cell. We also need to select a zinc particle distribution which could be easily dispensed and give good cell performance. The resulting mixture for the zinc anode paste is given in Table I.Table I : Zinc anode material formulation.Material% weight of zinc % weight of totalZinc Powder 8M aq. Potassium Hydroxide Gelling Agent100% 40% 2%70.4% 28.2% 1.4%296The gelling agent was selected based on recommendations from Norteman (1979), Kerg (1981), Bahary (1986), Graham et. al. (1987), Bennett et. al. (1997), and Sumiya et. al. (2003), which give information about starch graft copolymers, carboxylmethylcellulose, and polyacrylic acid polymers.

al. (1987), Bennett et. al. (1997), and Sumiya et. al. (2003), which give information about starch graft copolymers, carboxylmethylcellulose, and polyacrylic acid polymers. Multiple gelling agents from each of these groups were tested. The best results were obtained using a starch-based polymer with acrylonitrile side chains, commercially produced by B.F. Goodrich and known as Carbopol Ultrez 21. Graham and Goodman for Duracell Inc. (1987) note that “It has also been found that the use of (gelling agents of starch backbone and water soluble side chains) unexpectedly increases the practical discharge capacity of the anode while reducing the amount of cell gassing beyond that achieved through the use of previously known agents. Further, the internal resistance of the cell is not adversely increased.” Later research by Bennett et. al.

Further, the internal resistance of the cell is not adversely increased.” Later research by Bennett et. al. (1997) claims “the gelling agent acts to push zinc particles into spaces among its swollen particles to promote contact among the zinc particles, or between the zinc particles and the negative electrode current collector”.Table II : Performance of handmade cells with given amounts of gelling agent.Gelling Agent (wt% of zinc paste) Energy Output (J) Average Power Output (mW)0% 152.1 0.530.5% 193.8 0.541% 438.8 1.432% 768.4 2.25Cells with differing gelling agent concentrations were tested, as shown in Table II, and the outcome supports the claim that greater amounts of gelling agent increase cell capacity and power output. Gelling agent amounts in excess of 2% gave zinc anode pastes of a doughy consistency and were very difficult to extrude.

Gelling agent amounts in excess of 2% gave zinc anode pastes of a doughy consistency and were very difficult to extrude. Therefore a gelling agent amount of 2% was chosen as this gave the best output performance and provided reliable extrusion.Figure 1 : Zinc particle size distributions: (a) Standard Grade zinc powder (Big River Zinc), (b) Commercial Grade 2 zinc powder (Big River Zinc), (c) laboratory grade zinc powder (Sigma-Aldrich).The zinc particle size distribution was selected based on the recommendations of Chang et. al. (2003), Oyama et. al. (2004), and Durkot et. al. (2001), where the generally accepted particle sizes are between 50 and 325 mesh. Durkot states that the percentage of particles that pass through 200 mesh should be as much as 80 percent, and claims that cells with high proportions of fine zinc powder have increased performance especially in high drain usage. The patent offers the theory that this is due to particle-to-particle connectivity.

The patent offers the theory that this is due to particle-to-particle connectivity. Oyama states that both coarse and fine particles are needed, and that although increasing the amount of fine zinc particles improves discharge performance, it also increases generation of hydrogen gas through self-discharge. The Commercial Grade 2 distribution (Big River Zinc, Sauget IL) was chosen based on availability and ease of extrusion (Figure 1b). A greater percentage of fine particles,297passing through 200 mesh, permits the use of smaller nozzles without clogging. The Fine Powder distribution (Sigma-Aldrich, Inc.) has good extrusion performance; however, it displays significant hydrogen generation and causes problems with gas pockets in the extrusion syringes (Figure 1c).

The Standard Grade distribution (Big River Zinc, Sauget IL) requires a larger nozzle size to extrude reliably (Figure 1a) and tests with handmade cells indicate lower cell performance than with the chosen distribution (Table VII).Table III : Performance of handmade cells versus type of zinc powder employed.Zinc TypeGelling Agent (wt%) Energy (J) Ave. Power (mW) Continuous Service (h)Fine Powder 2 400.1 1.75 62.7Standard Grade 2 333.9 0.92 100.2Commercial Grade 2 2 768.3 2.29 94.4We prepare the zinc anode paste by first weighing the desired amounts of zinc, electrolyte, and gelling agent on a laboratory balance. We thoroughly mix the zinc powder and powdered gelling agent by hand to ensure that the gelling agent is evenly dispersed. We add the electrolyte and again mix thoroughly by hand until no electrolyte settles out of the mixture. The electrolyte concentration we employ is 8M which is typically used in the art.

We add the electrolyte and again mix thoroughly by hand until no electrolyte settles out of the mixture. The electrolyte concentration we employ is 8M which is typically used in the art. The percent weight of electrolyte in the anode is 40% which falls under the recommendations given by Chang et. al. (2003) (35-40%), Urry (2001) (25-55%), and Bahary (1986) (30-40%). Then we allow the paste to sit for at least 12 hours in order for the gelling agent to fully hydrolyze. When the paste has fully hydrolyzed no electrolyte should have separated out from the zinc. We mix the paste thoroughly again by hand and load it into its syringe.PVA SeparatorWe selected the materials for the separator based on the research presented by Treger et. al. (2003), who claim that a separator for an electrochemical cell can be formed in situ with a copolymer of polyvinyl alcohol and polyvinyl acetate. This technology proves particularly useful for application in freeform fabrication.

This technology proves particularly useful for application in freeform fabrication. This polyvinyl alcohol polymer, henceforth referred to as PVA, replaces the previous common technology of using unprintable paper (cellulose) separators. The PVA gel conforms to any geometry and when in contact with the electrolyte present in the zinc anode paste requires only 10 to 20 minutes to coagulate sufficiently to resume printing. The molecular weight distribution was changed slightly from the patent recommendation to favor a greater percentage of high molecular weight particles (Table VIII). We discovered that this change allows the separators to remain permanently flexible while not adversely affecting the cell performance. The gel is prepared by weighing the desired materials on a laboratory balance, mixing the dry materials, and slowly adding them to 60 degree water while stirring.

The gel is prepared by weighing the desired materials on a laboratory balance, mixing the dry materials, and slowly adding them to 60 degree water while stirring. The gel is mixed until the PVA crystals dissolve and then the gel is allowed to sit until the gas bubbles have escaped. The gel can then be loaded into the pistons for use.298Table IV: Polyvinyl alcohol (PVA) separator gel formulations.Material Polyvinyl alcohol (M.W. 31k - 50k) Polyvinyl alcohol (M.W. 85k - 146k) Polyvinyl alcohol (M.W. 124k - 186k) Deionized WaterOriginal formulation (Treger et. al., 2003)11% 3% 1% 84%Permanently flexible formulation10% 4% 2% 84%Catalyst Paste In creating the catalyst paste we need to adapt a typically dry mixture into one that could be extruded. Chang et. al. (2003) presents a catalyst formulation for metal air cells. We have adopted their recommendation for the ratios of dry materials (Table V).Table V. Catalyst dry mixtureTable VI.

Chang et. al. (2003) presents a catalyst formulation for metal air cells. We have adopted their recommendation for the ratios of dry materials (Table V).Table V. Catalyst dry mixtureTable VI. Catalyst paste formulationMaterial Manganese dioxide Oxidized Carbon Black PTFE 5 µm powderwt% 20 77 3Material Catalyst Dry MixtureCarbopol Ultrez 21 8M aq. KOHwt% 757 18To suit our purposes for extrusion and freeform fabrication we adapted this dry mixture by adding gelling agent and electrolyte (Table VI). The electrolyte serves the purpose of binding the dry mixture into a paste and the gelling agent prevents against phase separation. The electrolyte in the catalyst paste also assists in the coagulation of the separator layer. The paste is mixed by hand and then allowed to stand for twelve hours in order to fully hydrolyze.Case and Terminal MaterialsTwo additional materials are involved in the fabrication of batteries.

The paste is mixed by hand and then allowed to stand for twelve hours in order to fully hydrolyze.Case and Terminal MaterialsTwo additional materials are involved in the fabrication of batteries. A room-temperature vulcanizing (RTV) silicone (GE Silicone II, General Electric Silicones, Inc.) is used for the fabrication of the battery case and cap. This material is flexible, vulcanizes quickly (1-2 hours), is chemically resistant and inert, self-supporting, and easily deposited with 250µm resolution. The conductive terminals are produced using an RTV silicone filled with silver particles (SS-26, Silicone Solutions Inc.). This material is also chemically resistant and flexible, and bonds well to the silicone case to form a very robust conductor. The manufacturer claims an electrical resistivity of 5E-03Ω cm for the material.Freeform Fabrication ProcessThe batteries are produced using a multiple-material freeform fabrication system (Figure 2).

After the preparation of the materials, each material needs to be calibrated by varying the deposition control parameters until a uniform path of material is achieved. A three-dimensional CAD model of the desired battery is produced as an assembly. The model data is imported into the fabrication system software and material and fabrication properties are assigned to the parts of the assembly. Finally, the software generates a manufacturing plan from the design data, and executes the plan, using the syringe deposition tool and the various materials.299Figure 2. Schematic of our multiple-material freeform fabrication systemEach material allows a different degree of accuracy based on the smallest syringe tip that can be used to extrude reliably without clogging. The syringe tips are selected by hand extrusion testing with nozzles of decreasing diameter. The tips are disposable, and commercially produced by EFD Inc.

The syringe tips are selected by hand extrusion testing with nozzles of decreasing diameter. The tips are disposable, and commercially produced by EFD Inc. Stainless-steel needles are used for multi-phase materials, and plastic tapered nozzles are used for high-viscosity materials (Table VII). Tapered nozzles are very prone to clogging with suspended solid-phase materials; however, they allow high-viscosity materials to be extruded through small orifices without excessive pressure.

The first phase of parameter identification involves setting the material flow rate and positioning system traverse rate in order to achieve uniform and consistent paths of deposited material under steady state conditions. The second phase then involves adjusting the acceleration, deceleration, start delay, and stop delay of the material flow and the acceleration and deceleration of the positioning system in order to minimize the start/stop transient variations in path geometry. Finally, the height and width of the300deposited road are measured. We enter the parameters and road measurements into the materials database of the freeform fabrication system where it is accessed for automatic fabrication process planning.After the materials have been calibrated, the desired battery is modeled using a three dimensional mechanical CAD software package.

Each component of the battery is modeled as an individual part, and then parts are assembled to form a complete battery model (Table VI). We take into account the calibrated road geometry for the various materials during the design process in order to ensure that the design is feasible. All design features must be larger than the measured dimensions of the path for that material. The basic cylindrical cell has a diameter of 28.5 mm and a height of 10 mm (Figure 3).Figure 3. CAD model of standard cylindrical freeform fabricated battery; cutaway showing components (inset), and exterior view revealing cathode conductor, catalyst, and separator (cap hidden for clarity).We import the design geometry data into the system control software of our freeform fabrication system. The software graphical user interface (GUI) is used to assign a material from the materials database to each of the component parts of the battery.

The software graphical user interface (GUI) is used to assign a material from the materials database to each of the component parts of the battery. Each material is loaded into a separate disposable syringe with the appropriate tip. The syringes are installed into cartridges, which are then manually mounted in the deposition tool as requested by the system control software during the course of the build. We also use the GUI to assign a sequential fabrication priority to each of the parts to take advantage of what we refer to as “backfill deposition” (Figure 4b). Our SFF system will fabricate higher priority parts of an assembly prior to lower priority parts, in violation of strict layered fabrication. The priority will be obeyed by the system except where doing so would violate the relationship of one part supporting another.

The priority will be obeyed by the system except where doing so would violate the relationship of one part supporting another. Thus the silicone case of a battery can be given a higher priority than the materials to be deposited into it, and it will be completely fabricated to its full height before the deposition of any other materials into it begins. Backfill deposition allows the case to act as a container for materials which are not self- supporting, greatly improves the quality of deposition, and drastically reduces the number of material changes required which decreases build time.301When fabricating a standard cylindrical battery, we set the battery case and the anode conductor to the same priority so the conductor can be build through an opening in the wall of the case for ease of connection (Figure 4a). The case is completed and the zinc anode layer and PVA separator layer are deposited using backfill (Figure 4b).

The case is completed and the zinc anode layer and PVA separator layer are deposited using backfill (Figure 4b). The build is paused for 15 minutes to allow the PVA to react with the electrolyte and coagulate so that the separator will not be damaged by the remaining layers (Figure 4c). The catalyst, cathode current conductor, and casing cap layers are then deposited to complete the battery (Figure 4d). A standard cylindrical battery requires roughly 1.25 hours to be produced. In order to monitor the mass of active materials that is deposited into the battery, we mount a laboratory balance atop the build surface, and fabricate the battery onto a sheet of polyester film attached to the balance.Figure 4.

When the cells fall below 0.1 volts we stop the recording and remove the batteries.Results and DiscussionThe resultant discharge curves of the freeform fabricated batteries are shown in Figure 5. The curves are characteristic of zinc air batteries which have a long period of constant voltage followed by a rapid drop-off. As expected the triple cell area has the highest voltage and operating current, however it has the shortest lifetime. The standard cell shows a slightly lower voltage and operating current, and a slightly longer lifetime than the triple-area cell. The half area cell has a much lower voltage and operating current, however it achieves a lifetime twice that of the standard cell.302Based on these results, we can estimate the specific current capacity (Figure 6), service lifetime, and average current (Figure 7) of a freeform fabricated battery of a given intermediate active surface area.

A smaller surface area gives a lower current because there is less area for reactions to take place. As the surface area increases the current reaches a limiting value for a given load. The open circuit potential for a zinc-air cell is about 1.4V, though the nominal loaded voltage drops to about 1.2V and higher current densities result in even lower output voltages. Mao and White (1992) explain the drop to 1.2V as the initial depletion of free OH- ions from the electrolyte, and the larger drop with increased current density as reaction-rate limitation. In our testing, cells with greater surface area achieved higher operating voltages, although none achieved the 1.2V nominal voltage. This suggests that we are operating at too high of a current density for the design or materials employed in our cells. In addition, the greater the exposed surface area the shorter the lifetime, implying that dehydration may also be a limiting factor.Figure 5.

In addition, the greater the exposed surface area the shorter the lifetime, implying that dehydration may also be a limiting factor.Figure 5. Discharge curves for cylindrical batteries of three different active surface areas.303Figure 6. Specific current capacity (to mass of zinc in anode) of cells versus their active surface area. Error bars are 1 standard deviation for sample size of 2.Figure 7.Average current from cells over service life into 100 Ohm load versus surface area.304Table VIII : Comparison of commercial, handmade, and freeform fabricated cells.CommercialFreeform FabricatedHandmadeOpen Circuit Voltage (V)Cont. Service Life (h)Eveready AC675 (2004) 1.40 302 (@16h/day)Basic 3 Basic 4 Triple 1 Half 11.381.091.291.2927.4830.0628.2063.00HM11.3338.90HM21.3469.61Fabrication Time (h)1.251.431.83Energy Output (J)2996.6416.59570.63736.07543.55760.27576.00Ave.

Power (mW)2.764.285.287.252.413.562.29Ave. Current (mA)2.16.427.258.484.875.964.72Capacity to .4V (mAh)635176.45217.83239.27306.88231.80328.56Mass of Anode (g)~ 14.994.996.365.123.3022.90Mass of Zinc (g)<13.513.514.483.616.4016.12Surface Area (cm2)1.054.684.6814.042.345.075.07Specific Energy (kJ/kg)1577.16118.59162.44164.40151.09157.74145.80Ave. Power/Area (mW/ cm2)2.630.911.130.521.030.771.03Table VIII gives a full listing of all result data and calculations. Data from zinc air button cells is listed for comparison. The comparison to commercial cells is not accurate, in that commercial batteries are tested with a load of 625Ω and therefore discharge at a lower current density. This may explain some of the disparity in performance.

This may explain some of the disparity in performance. Improvements to cell capacity will include a humidity control layer, such as PTFE dispersion, design changes to reduce electrolyte evaporation, and adjustments to the separator layer to prevent the zinc anode from becoming exposed.To fully utilize the possibilities of freeform fabrication, we have designed and produced a flexible, two-cell, zinc air battery with an unusual geometry. We chose a U-shape where the layers were organized horizontally instead of vertically to allow air flow to both catalyst layers, as shown in the CAD model (Figure 8a). The completed cell measures 11 centimeters long and 8 centimeters wide (Figure 8b). The cell provided an open circuit voltage of 2.5 V, clearly indicating the functioning of both cells in the two-cell configuration. After overnight testing with a 100Ω load, the cell still produced an open circuit voltage of 1.41 V.

After overnight testing with a 100Ω load, the cell still produced an open circuit voltage of 1.41 V. The loss is likely due to breakage in the separator layers which allows internal shorting. Increasing the thickness of the separator layers should eliminate the internal shorting. The geometry of the battery and the material formulations allow the battery to be flexed (Figure 8c) and the battery maintains the 1.4V output even when flexed (Figure 8d). This demonstration shows only one example of the geometries that are possible using a freeform fabrication system.305Figure 8.

This demonstration shows only one example of the geometries that are possible using a freeform fabrication system.305Figure 8. (a) Labeled CAD model of two cell U-shaped battery, (b) completed battery, (c) demonstration of the flexibility of the battery, (d) voltage after testing.ConclusionThrough an extensive set of experiments with handmade zinc-air batteries, and taking advantage of the commercial research represented by numerous patents on traditionally manufactured batteries, we have developed new formulations for the zinc anode, cathode catalyst, and separator materials which are suitable for freeform fabrication. We have produced and tested a flexible, two-cell battery with an unusual shape that shows the type of geometric and functional customization that freeform fabrication brings to the manufacture of batteries. We have collected performance data for a series of freeform fabricated batteries of identical design but varied diameter.

We have collected performance data for a series of freeform fabricated batteries of identical design but varied diameter. This data allows the estimation of the performance of a battery of a similar proposed design prior to fabrication. We are continuing to collect data from batteries of widely varying designs to more fully identify the effects of geometry on freeform fabricated battery performance. Using our data for freeform fabricated batteries, we will extend the predictive model of Mao and White (1992) to cover more complex geometries and multi-cell batteries. This will lead to a true predictive performance model which can be used to design geometrically and performance-optimized zinc-air batteries for incorporation into many different electrical applications.306ReferencesAkashi, H., K. Tanaka, et al. (2002). "A flexible Li polymer primary cell with a novel gel electrolyte based onpoly(acrylonitrile)." Journal of Power Sources 104(2): 241.Bahary, W. S. (1984).

EEFORM FABRICATION OF ZIRCONIUM DIBORIDE PARTS USING SELECTIVE LASER SINTERINGMing C. Leu1, Erik B. Adamek1, Tieshu Huang2, Greg E. Hilmas2 and Fatih Dogan2 1Department of Mechanical and Aerospace Engineering 2Department of Material Science and EngineeringMissouri University of Science and TechnologyRolla, MO 65409{mleu@mst.edu, ebawc3@mst.edu, hts@mst.edu, ghilmas@mst.edu, doganf@mst.edu}Reviewed, accepted September 10, 2008ABSTRACTUsing the Selective Laser Sintering (SLS) process, both flexural test bars and 3Dfuel injector components have been fabricated with zirconium diboride (ZrB2) powder. Stearic acid was selected as the binder. Values of SLS process parameters were chosen such that the green parts could be built with sharp geometrical features and that the sintered parts could have good mechanical properties.

Values of SLS process parameters were chosen such that the green parts could be built with sharp geometrical features and that the sintered parts could have good mechanical properties. After binder burnout and sintering, the SLS fabricated ZrB2 test bars achieved 80% theoretical density, and the average flexural strength of the sintered samples was 195 MPa. These values demonstrate the feasibility of the SLS process for freeform fabrication of 3D parts with the ultra high temperature ceramic.1. INTRODUCTIONComplex three-dimensional shapes are not always a viable option from ceramic materials using conventional manufacturing techniques. In addition to high failure rates in machining operations such as drilling and milling due to brittleness of ceramic materials, some geometric features like internal voids and curved holes are impossible to create by these and other conventional methods.

Although these techniques have been demonstrated for layer-by-layer freeform1fabrication of ceramic components, many technical challenges remain, especially for SFF of ultra high temperature ceramics (UHTCs).Selective Laser Sintering (SLS) holds two unique advantages compared withother SFF processes: one is the wide variety of materials that can be used in SLS processing, and the other is no need for building, and subsequently removing, support structures in fabricating complex 3D parts. In the SLS process, the STL file of a solid model is sliced into layers of specified thickness. The contours of each layer are transferred to the process controller of an SLS machine, which actuates servo mirrors to direct a laser beam to outline and fill in the two dimensional profile of a layer. After each layer is completed, the build area lowers by one layer thickness and fresh powder is spread over the build area. The layer fabrication step repeats until the entire part has been built.

After each layer is completed, the build area lowers by one layer thickness and fresh powder is spread over the build area. The layer fabrication step repeats until the entire part has been built. After finishing the build, the part cake is removed and broken to obtain the built part. In SLS of ceramics, the fabricated green part is further processed for binder burnout and subsequent sintering to achieve high relative density and mechanical strength. A comprehensive discussion of consolidation phenomena in SLS (and its variants) of various types of materials is given by Kruth et al. [2007].Zirconium Diboride (ZrB2) is an ultra high temperature ceramic with a melting temperature of 3245 °C, a flexural strength of 200-590 MPa, and a hardness of 21-23 GPa [Fahrenholtz et al., 2007]. It has high electrical and thermal conductivity as well as excellent chemical resistance and thermal shock resistance [Bellosi and Monteverde, 2002; Fahrenholtz et al., 2007].

It has high electrical and thermal conductivity as well as excellent chemical resistance and thermal shock resistance [Bellosi and Monteverde, 2002; Fahrenholtz et al., 2007]. ZrB2 is an excellent candidate for a number of ultra high temperature applications, such as rocket propulsion nozzle throat inserts and leading edges of future-generation hypersonic space vehicles. However, since ZrB2 cannot be cast using traditional forming processes, due to its extremely high melting temperature, current processing techniques are limited to pressure-assisted and pressureless sintering followed by post-process machining. Not only the brittleness of material limits the scope and success rate of the post-process operations, but the rate of manufacture is necessarily low in order to reduce the chances of failure.Previous research on SLS of ZrB2 powder included a study by Stucker et al. [1995, 1997] to develop a new way of manufacturing electrical discharge machining (EDM) electrodes.

[1995, 1997] to develop a new way of manufacturing electrical discharge machining (EDM) electrodes. ZrB2 green parts were fabricated by SLS, followed by debinding, sintering, and infiltration with copper to make EDM electrodes. The sintered ZrB2 parts were 31% dense on average before Cu infiltration. There have been no mechanical properties reported for these sintered parts. Another study of laser sintering of ZrB2was performed by Sun and Gupta [2008], who reported that ZrB2 powder having small particle size improved surface wetability and sinterability.The present paper describes an experimental study of using the SLS process to fabricate ZrB2 components. The experiments involved consideration of materials, SLS process parameters, binder burnout, and sintering. The process was developed to fabricate ZrB2 samples including flexural test bars and 3D fuel injector components. The dimensions of both green and sintered test bars were measured and compared with nominal dimensions.

The dimensions of both green and sintered test bars were measured and compared with nominal dimensions. The measured dimensions, density and flexural strength of the test samples are presented and discussed.22. EXPERIMENTS2.1 Materials and Powder PreparationThe material used was zirconium diboride powder (grade B, H.C. Starck, Karlsruhe, Germany) with an average particle size of 2-3 µm and a relative surface area of 1.38 m2/g (8.42 m2/cm3). The as-received ZrB2 powder had an oxygen content of 0.9 wt.%. Boron carbide (B4C; grade HS, H.C. Starck) having an average particle size of 0.8 µm and a surface area of 15.8 m2/g (39.82 m2/cm3) was used as a sintering additive. The oxygen content of the B4C additive was about 1.3 wt.%. Carbon black (Black pearls 120, Cabot Corporation, Alpharetta, GA) was used as a source of carbon (C).

The oxygen content of the B4C additive was about 1.3 wt.%. Carbon black (Black pearls 120, Cabot Corporation, Alpharetta, GA) was used as a source of carbon (C). Stearic acid (Acros Organics, Stearic Acid, 97% New Jersey) was used as the binder.To prepare ZrB2 powder for use in the SLS process, first the powder was mixed with sintering additives (1 wt.% B4C and 0.2 wt.% C) by ball milling for a period of 24 hours. These additives promote full sintering during the post-SLS sintering process. After the ceramic and additives were mixed thoroughly, stearic acid was added at 50:50 vol.% and ball milled for an additional period of at least 24 hours. Stearic acid was chosen as the binder for the ZrB2 powder based on the result of a previous study performed for SLS of Al2O3 powder [Liu et al., 2007]. Stearic Acid has a number of beneficial properties for use as the binder. It is a thermoplastic and the change in its state from a solid to a liquid yields only a small change in density.

Stearic Acid has a number of beneficial properties for use as the binder. It is a thermoplastic and the change in its state from a solid to a liquid yields only a small change in density. This reduces curling in the part bed and contributes to more successful lamination. Stearic acid does not react with ZrB2 and the powder can be reused several times. The low melting temperature of the binder adds to its SLS friendliness. The 50:50 ZrB2/stearic acid mixture was used for the SLS process after the initial 55:45 ZrB2/stearic acid mixture yielded no successful fabrication.The mixed powder was then checked for clumping, which may occur if the humidity and temperature are high due to the waxy nature of stearic acid. Problems faced during powder processing included some batches of powder agglomerating into a ball or a ring of powder as shown in Figure 1. These clumps were crushed and re-milled to return to powder form.

These clumps were crushed and re-milled to return to powder form. The differences between batches were the ambient humidity and temperature. The moisture content has the largest effect, causing the powder to clump together perchance during SLS processing.(a)(b)Figure 1: Clumps of powder (a) ball and (b) ring32.2 SLS Processing and Process ParametersThe SLS system used was a DTM 2000 Sinterstation shown in Figure 2. Core features of this machine include heated powder feed bins as well as heating of the part piston. A 50W CO2 laser was used to introduce beam energy into the part bed by selectively directing laser beam onto the surface of the part bed.

A 50W CO2 laser was used to introduce beam energy into the part bed by selectively directing laser beam onto the surface of the part bed. Preparation of the setup for SLS processing involved placing .STL files for the parts into the Sinterstation software package, setting all process parameters, filling the feed bins with powder, and sealing and evacuating the O2.The values of SLS process parameters were selected by a combination of theoretical calculations, analytical reasoning, and empirical testing. The temperature of the powder feed bins must be low enough so that the binder is not sticky while high enough so that it can heat the powder quickly in the build area. The part piston and part bed temperatures need to be close to the melting temperature of stearic acid (69°C) so that the laser does not need to excite the powder temperature to such a degree that curling due to the change in density as a result of temperature gradient becomes inevitable.

A temperature that is too high, however, may result in one fused piece. The layer thickness at 0.0762 mm is the smallest layer thickness available from the Sinterstation. This thickness was chosen because the layers did not laminate together when the thickness was set any higher, due to the limited penetration ability of laser in ZrB2 powder.The laser power (W) is related to the process parameters as follows: ][ ()kWScanSpeed×H+×ScanSpacin g×Hp 1T×Δ×=Hp 221 cf2c×LayerThickwhere TΔ is the change in temperature, cH is the heat capacity of material x, x,xp is the volumetric percentage of material fH is the heat of fusion of the binder, and k is a coefficient depending on powder density, reflectivity, etc. A series of experiments were conducted to determine the scan speed and laser power.

A series of experiments were conducted to determine the scan speed and laser power. The experiments indicated that the laser has limited penetration ability in ZrB2 powder, thus the laser scan speed has to be decreased substantially from that previously used in SLS of Al2O3 [Liu et al., 2007] in order to generate sufficient binder fusion and avoid delamination between layers. The amount of energy required was determined assuming perfect absorption by the powder and no significant heat dissipation, thus the calculation serves as a low estimate of the necessary power setting and is a first approximation.Based on the above considerations, the following SLS process parameters werechosen for fabricating ZrB2 parts: part bed temperature 59°C, part piston temperature 48°C, feed bin temperature 25°C, laser scan speed 50.8 mm/s, scan spacing 0.152 mm, layer thickness 0.0762 mm, and laser power from 0.8 to 1.0 watt.

With this set of parameters the SLS process yielded 15 flexure test bars and 7 injector parts, representing an overall yield of 17%.ness4Laser WindowRollerBuild AreaPart PistonPowder BinFigure 2: DTM 2000 SinterstationZrB2 parts fabricated by the Sinterstation were broken out using a combination ofmechanical brushing for rough breakout and compressed air around 10 psi for fine surface cleaning. The green parts were inspected visually for possible delamination and fracture during the fabrication process. Figure 3 shows the ZrB2 powder bed with 12 fused parts during the SLS processing. The absence of curling and the change in color due to the vaporization of stearic acid indicates that the SLS process was going smoothly. However, the quality of lamination between layers cannot be judged by visual observation during the fabrication process. The acceptability of layer-to-layer lamination can only be determined after part breakout.

The acceptability of layer-to-layer lamination can only be determined after part breakout. Figure 4 shows the sharp features of green ZrB2 test bars after SLS processing.Figure 3: ZrB2 test samples in the powder bed during SLS processingFigure 4: Green ZrB2 test bars after SLS processingDoor52.5 Binder Burnout and SinteringBinder burnout for the green ZrB2 parts was performed in a furnace using thetemperature profile shown in Figure 5. The parts were slowly brought to a temperature of 600°C. The ramp time for reaching this temperature was chosen specifically to be very long in order to allow vacating the binder completely from the part without bloating or cracking. The process took about 52 hours. It removed all of the organic species and allowed for some rearrangement of the ceramic particles (resulting in a minimal amount of shrinkage). The binder burnout for ZrB2 samples was accomplished in an environment of 10% H2 and 90% Ar.

The binder burnout for ZrB2 samples was accomplished in an environment of 10% H2 and 90% Ar. This environment was used to remove the binder without unwanted oxidation of the ZrB2 powder. The binder burnout process had a 100% yield.The green test bars underwent isostatic pressing before the binder burnout. Theisostatic pressing was done by vacating the air in a nitrile glove with parts located at the fingers and dipping the glove into the isostatic pressing bath at 276 GPa for 1 min. It was used to increase the green densities of the parts in order that higher sintering densities can be achieved.

It was used to increase the green densities of the parts in order that higher sintering densities can be achieved. This process had a 100% yield.600)C ° (500e r u t a r e p m e T40030020010000102030 Time (h)405060Figure 5: Binder burnout schedule for ZrB2 samples2500)C ° (2000e r u t a r e p m e T1500100050000100200300400500600700Time (min)Figure 6: Sintering schedule for ZrB2 samples6Final sintering of ZrB2 parts was performed at 2,200°C, which was maintained for2 hours in a Lindberg graphite furnace (model 51644). The parts were then cooled at a rate of 10 ˚C/min. The sintering schedule is shown in Figure 6. The sintering process took about 10 hours, and it had a 100% yield.3. RESULTSTo study the physical properties and microstructures of SLS processed ZrB2 parts,flexural test bars and 3D fuel injector components were fabricated using the Sinterstation.

RESULTSTo study the physical properties and microstructures of SLS processed ZrB2 parts,flexural test bars and 3D fuel injector components were fabricated using the Sinterstation. The green parts then underwent binder burnout and sintering. Figure 7 shows a green ZrB2 test bar and the same bar after sintering. Figure 8 shows a fuel injector part viewed in two different directions and the same part after sintering.(a)(b)Figure 7: A ZrB2 flexural test bar made by SLS: (a) green part and (b) after sintering. The dimensions are: for the green part, X=51.8 mm, Y=6.8 mm, Z=6.2 mm; for the sintered part, X=42.0 mm, Y=5.2 mm, Z=3.7 mm.Figure 8: A ZrB2 3D test part made by SLS: (a) green part, (b) another view of the test part, and (c) after sintering.

(mm) 51.9 52.0 52.0 52.0 52.0 52.0 51.9 51.9 51.9 51.8 51.9 51.9 51.9 51.8 51.92Green Y Dim. (mm) 6.9 6.8 7.0 6.8 6.9 6.8 6.8 6.8 6.8 6.8 6.8 6.8 6.8 6.8 6.83Sintered Y X Z Dim. Dim. Dim. (mm) (mm) (mm) 5.4 42.2 6.7 5.3 42.3 6.8 5.3 42.7 7.0 5.3 42.1 6.5 5.3 42.4 6.2 5.3 42.2 6.7 5.1 41.9 6.6 5.1 42.1 6.7 5.2 42.0 6.4 5.2 42.0 6.2 5.2 42.0 6.2 5.2 42.3 6.6 5.2 42.3 6.3 6.7 5.2 42.1 6.54 42.19 5.24Standard Deviation0.080.060.240.220.08Table 2: Dimensions of the green and sintered test bars as percent of nominal dimensionsXYZGreen Part% of nominalstd. deviation (% of nominal)1000.161030.90973.59Sintered Part% of nominalstd. deviation (% of nominal)810.41791.25594.30The manufactured parts were measured for both the green and sintered parts.Table 1 shows the dimensions measured from the 14 successfully sintered test bars against nominal dimensions.

Table 2 shows the dimensions of the green and sintered test bars in percentage of the nominal dimensions. Note that X was along the scanning direction, Y was perpendicular to the scanning direction, and Z was in the depth (vertical) direction. All test bars were fabricated so that the longest dimension was the direction of laser travel (i.e. X direction). The data in these two tables indicates that the X dimension measured from the green test bars was only 0.1% difference on average from the nominalZ Dim. (mm) 4.4 4.3 4.6 4.0 3.9 4.2 3.8 3.8 3.7 3.7 3.7 3.9 3.6 3.9 3.960.298dimension, and that the Y and Z dimensions measured from these test bars were less than 4% different from the nominal dimensions on average. After sintering, the part shrank to about 80% of the nominal dimensions in the scanning plane (both along and perpendicular to the scanning direction) and about 60% of the nominal dimension in the depth direction.

These results indicate that there is significantly more fusion in the laser scanning plane and less fusion in the depth direction. Consequently, more shrinkage occurred in the depth direction during sintering.After sintering, the densities of ZrB2 test bars were measured using Archimedes method. The average density measured of the sintered test bars was 4.87 g/cc, which is 80% theoretical density. This is much higher than the 31% theoretical density previously reported [Stucker, 1997]. The density of sintered part is affected by the ceramic/binder ratio in the powder mixture and the ZrB2 particle size. In our study, we had the 50:50 ratio of ZrB2/stearic acid in the powder mixture. The ZrB2 powder as received had ∼3 µm in particle size, and the particles were not attrition milled. Attrition milling can significantly reduce the particle size and increase the sintering drive force for densification because smaller particles provide more surface areas overall.

Attrition milling can significantly reduce the particle size and increase the sintering drive force for densification because smaller particles provide more surface areas overall. A larger surface area is associated with a higher surface energy, which is the major driving force for pressureless sintering.Figure 9 shows the microstructures from a fracture surface of a test bar in twodifferent magnifications. It can be seen that the microstructure is fairly uniform and that the pores are on the order of several microns to about ten microns in size. The fracture path is predominantly transgranular, indicating that the fracture is a typical type of brittle fracture as expected.Figure 9: Microstructures of a test bar in two different magnificationsThe flexural strength was measured from the 14 sintered test bars using the fourpoint bending method.

All test bars were ground to “A” bar size per ASTM standard C1161 “Standard Test Method for Flexural Strength of Advanced Ceramics at Ambient Temperature.” This standard has the test bar’s cross-sectional area of 2.0 mm by 1.5 mm (±0.05mm). By using an Instron machine, the flexural strength data was measured and collected as given in Table 3. The mean flexural strength measured from these test bars was 195 MPa, which is 44% of the flexural strength of 444 MPa obtained from hot9pressed ZrB2 [Chamberlain et al., 2006]. To our knowledge this is the first published result on the flexural strength of SLS fabricated ZrB2 test bars. The relatively low strength of these test bars compared to that of hot pressed samples was associated with its relatively high porosity. The hot pressed samples after sintering had 98% theoretical density, while the SLS processed samples after sintering had 80% theoretical density.

The hot pressed samples after sintering had 98% theoretical density, while the SLS processed samples after sintering had 80% theoretical density. Nevertheless, the 80% theoretical density achieved in our study is much higher than the 31% theoretical density previously reported [Stucker 1997]. In order to further increase the density, our future work will include optimizing the powder/binder compositions and SLS process parameters.Table 3: Flexure strength of the test barsPart No.1 2 3 4 5 6 7 8 9 10 11 12 13 14 Mean Standard Deviation Hot Pressed ZrB2Flexure Strength (MPa) 218 216 208 205 205 200 199 197 190 185 184 183 178 162 195 16 4444.

CONCLUSIONSLS processing of ZrB2 powder with stearic acid as the binder has been studied.Flexural test bars and 3D fuel injector parts have been successfully fabricated by the SLS process using a zirconium diboride/stearic acid mixture of 50:50 vol.% with appropriately determined values for SLS process parameters including part bed and feed bin temperatures, laser scan speed, scan spacing, layer thickness, and laser power. After binder burnout and sintering, the flexural test bars achieved 80% theoretical density, which is much higher than the density of similar test bars previously reported in the literature. The average flexural strength of the sintered test bars was 195 MPa, which is 44% of the flexural strength of hot pressed ZrB2 test bars.

The average flexural strength of the sintered test bars was 195 MPa, which is 44% of the flexural strength of hot pressed ZrB2 test bars. The SEM images from the fracture surface of the samples indicated that the microstructure was fairly uniform and the pore size was about several to ten microns.10ACKNOWLEDGMENTSThis work was supported by the Air Force Research Laboratory under Contract FA8650- 04-C-5704.REFERENCES1 Bandyopadhyay, A., Panda, P. K., Agarwala, M. K., Danforth, S. C., and Safari, A., (2000) “Processing of Piezocomposites by Fused Deposition Technique,” Journal of American Ceramic Society, 80(6): 1366-72.2 Beaman, J. J., Barlow, J. W., Bourell, D. L., Crawford, R. H., Marcus, H. L., andMcAlea, K.

EEFORM POWDER MOLDING FOR RAPID TOOLING Stephen J. Rockt+, Charles R. Gilmant +, Wojciech Z. Misiolekt , and Daniel F. Walczykt tRensselaer Polytechnic Institute New York State Center for Advanced Technology in Automation, Robotics and Manufacturing Troy, New York 12180 -and- +Manufacturing Solutions, Inc. Troy, New York 12180ABSTRACTTooling development can be quite time consuming and costly. Several iterations may be required, and later product or process modifications may necessitate tooling redesign. Rapid prototyping techniques capable of meeting the structural requirements of short-run and end-use tooling will have a significant impact on the product development cycle. This paper presents a technique for producing tooling using the Freeform Powder Molding process. Resulting tooling can be made from a wide variety of readily available metal powders, and mechanical properties can be tailored for customized tool design and fabrication.

Resulting tooling can be made from a wide variety of readily available metal powders, and mechanical properties can be tailored for customized tool design and fabrication. The example presented in this paper focuses on the rapid production of tooling for sheet metal forming.1. INTRODUCTIONSince the emergence of the first commercial Solid Freeform Fabrication (SFF) processes, SFF system production rates have been unable to match the rates of conventional high-volume manufacturing operations such as injection molding, die casting, and stamping. While substantial progress has been made toward producing SFF components from engineering materials, throughput limitations continue to remain a formidable obstacle to the deployment of SFF systems in production.

If these systems can be used to produce tooling for high-volume manufacturing processes, then throughput limitations can be surmounted while realizing many of the benefits associated with SFF, and more generally with Rapid Prototyping (RP). This use of RP to make tooling offers one means for realizing Rapid Tooling (RT). Rapid tooling allows early implementation of production manufacturing techniques and materials to evaluate product designs earlier in the development cycle than is possible where conventional techniques are used to fabricate tooling.This paper introduces a technique for quickly producing tooling based on the two-material Freeform Powder Molding (FPM) concept [1], and the notion of freeze molding [2]. While much work has been published on the use of RP parts as patterns [3-5], particularly for casting applications [6-8], this work describes the direct use of an RP part as a mold for shaping metal powder.

The resulting molded metal object then serves as tooling for sheet metal forming. Before discussing details of tooling production, a brief introduction to FPM and freeze molding are provided.1.1 Freeform Powder MoldingThe FPM process relies on material behavior differences to selectively define component shape [1]. By arranging powders with different diffusion kinetics within some confining volume, it is possible to process the entire volume and cause only one of the materials to consolidate while the other remains a loose powder. This is illustrated in Figure 1, where material that consolidates is termed part powder and material which does not consolidate is termed tool powder. In this example, the part powder represents a part which will become the tooling used for metal forming.149PartPowder~~~. . . ..

In this example, the part powder represents a part which will become the tooling used for metal forming.149PartPowder~~~. . . . . .Tool PowderPowder MassResulting PartFigure 1 - Freeforrn Powder Molding Two Material Concept [1]The tool powder essentially forms soft tooling which both supports, and in conjunction with the part powder, defines the shape of the resulting component.Although originally envisioned as a layer-wise SFF process, the underlying process concept can be applied when alternative shape defining techniques such as freeze molding are employed. Powder can be shaped by molding, and then surrounded with tool powder to preserve its shape during carrier melting and subsequent removal.1.2 Freeze MoldingFreeze molding, also referred to as freeze casting, presents a low-cost alternative to powder processing techniques requiring costly tool sets.

Freeze molding was originally explored to bring concurrency to FPM process development and distortion compensation research [9]; however, the results of this work suggest it has merit as a secondary material conversion alternative for RP. A slurry of powder and carrier is introduced into a mold at low (or no) pressure and the mixture is frozen [2]. Molds can be fabricated from a variety of inexpensive materials which can be easily machined or they can be produced by a variety of commercial RP processes. Since the powder slurry is not injected into the molds under high pressure, the geometric complexity of the molds may be limited.Aqueous carriers are used because they can be easily frozen. Carriers may contain additional constituents which support the development of green strength after shaping of the powder compact and sublimation of the principal carrier component.

Carriers may contain additional constituents which support the development of green strength after shaping of the powder compact and sublimation of the principal carrier component. The presence of some organic binder is particularly important if the compact is dried rapidly, as this can cause cracks within the compact [10]. Unfortunately, carrier sublimation takes a long time and this disadvantage is similar to the debinding operation necessary with powder injection molding [2].While typical freeze molding requires carrier sublimation and compact green strengthdevelopment, these are not required when coupling freeze molding with FPM.2. RAPID TOOLING PRODUCTION PROCEDUREThe following procedure is used to produce tooling. A mold is created by RP and used to shape a powder mass. This powder is subsequently sintered to realize a solid component which has the physical properties required of tooling.

A mold is created by RP and used to shape a powder mass. This powder is subsequently sintered to realize a solid component which has the physical properties required of tooling. The resulting tooling is used to make the final desired component, which in this application is formed sheet metal. Figure 2 depicts the sequence of operations used for rapid tooling production. While oversimplified by this depiction, this procedure was quite straight-forward and efficient to implement. The complete rapid tooling sequence from die design to sheet metal forming could be completed in only four days given this simple forming geometry.2.1 Powder Selection & PreparationRapid tooling was fabricated using both Anval 316 stainless steel powder with a particle size of -325 mesh/+ 16 Jlm and Novamet 4SP elemental nickel powder with a particle size of -20 Jlrn/+ 10 Jlm. Both powders are chemically compatible with the carrier used during freeze150moldin2:.

Both powders are chemically compatible with the carrier used during freeze150moldin2:. Both powders also have relatively small particles to promote rapid sintering to maximum density~ given loose powder preforms. The small parti~les help to dev~lop a fine. grained microstructure in the resulting sinter-body which leads to Improved. me~h.a~Ical pr?pertles.. Both powders have particles of spherical morphology so they pac~ to a hIgh InItIal densIty, relative to irregularly shaped powder, in the absence of external compaction pressure.Die DesignMold DesignRP MoldSintered DieFormed PartFigure 2 - FPM Rapid Tooling ProcedureAn aqueous carrier comprized of water with 10% by volume ethyl alcohol to facilitate particle wetting was used in this work. This carrier was selected because it was easily frozen and later removed from the powder.

This carrier was selected because it was easily frozen and later removed from the powder. No additional organic binding constituents were necessary since the modified freeze molding procedure does not require that the compact develop green strength or that it withstand sublimation which may lead to crack formation. Each powder was manually stirred into the carrier, and sufficient carrier was used to ensure the powder remained completely saturated.The powder used to support the rapid tooling in the frozen state and during sintering, known as tool powder given the FPM process definition, was Norton #204 yttria zirconia. Also spherical in shape, this powder has a broad powder size distribution below 180 /..lm.

Also spherical in shape, this powder has a broad powder size distribution below 180 /..lm. It is important that this powder have sufficiently different diffusion kinetics such that it will not sinter at the given processing conditions.2.2 Mold Fabrication & FillingMolds suitable for shaping the powder slurry described previously can be made in a variety of ways provided the mold material can withstand the temperatures required to freeze the powder slurry and is compatible with slurry chemistry. To realize the many benefits of rapid prototyping which contribute to this rapid tooling process, the molds should be produced using manufacturing automation guided by CAD model information. A major goal of this effort was to show that RP parts can be used directly to cast the powder slurry. Molds were produced both by the Stereo lithography process and by NC machining. These are shown in Figures 3 and 4, respectively.

Molds were produced both by the Stereo lithography process and by NC machining. These are shown in Figures 3 and 4, respectively. Relatively simple die shapes were selected for the sheet metal forming example in this paper to simplify mold design and fabrication.Figure 3 - SOMaS 6110™ SLA Mold Set151Figure 4 - Acetal Resin NC Machined Mold SetThis tooling geometry is similar to the ball punch deformation test specified by ASTM E643 [11]. Sintering furnace size also restricted the tooling size.The molds were designed to produce dies with a suitable clearance to accommodate the 0.033" sheet metal thickness. A draft angle of 3° was designed into the vertical walls of the mold to facilitate part ejection. Each mold is designed to attach to a backing plate (not shown) which provides a flat surface on the back of each metal forming tool necessary for attachment to press equipment. Molds are filled from one side via a rectangular gate.

Molds are filled from one side via a rectangular gate. Mold filling is performed manually under the influence of gravity. The molds are lightly tapped as the powder slurry is added to facilitate the powder fill. 2.2.1 Stereolithography MoldsOne set of molds was fabricated directly by Stereolithography (SLA) using a 3D Systems SLA-250/40. The molds were fabricated from the recently introduced DuPont SOMOSTM 6110 epoxy photopolymer resin using 0.006" thick build layers.Geometry for these molds was designed using Pro/Engineer and exported as an STL facet file with a chordal deviation set to 0.0004" to minimize tessellated surface artifacts. Such artifacts were nearly impossible to observe in the manufactured parts; however, the well-documented spatial aliasing (stair-stepping) caused by layer-wise fabrication was evident. The molds were used in as manufactured condition without manual post-processing (e.g.

The molds were used in as manufactured condition without manual post-processing (e.g. sanding, polishing, etc.), with exception of the manual filling of a small hole added for trapped-volume pressure equalization during resin recoating [12]. Prior to slurry filling a silicone-based molding release agent was applied to the SOMOS 6110 material.2.2.2 NC Machined MoldsAnother set of molds was fabricated by NC machining using an acetal resin (DuPont Delrin™). This material machines easily and yields a high quality surface finish without subsequent hand finishing. It has been used extensively in other freeze molding shaping applications [9].Molds were fabricated by NC machining to illustrate the potential role that more conventional manufacturing processes hold for rapid tooling. The benefits of solid modeling and ease with which information is handled by SFF processes became painfully clear during the 2D mold design and CAM path planning steps leading to cutting the molds.

The benefits of solid modeling and ease with which information is handled by SFF processes became painfully clear during the 2D mold design and CAM path planning steps leading to cutting the molds. These activities accounted for much greater delay than actually cutting the plastic mold material. However, for a fair comparison to be made between SFF and NC machining RP approaches, both processes should be driven from a 3D solid model.2.3 Modified Freeze MoldingA variant of the freeze molding procedure described in section 1.2 was used to shape the powder which ultimately becomes the rapid tooling of interest. Molds filled with the powder/carrier slurry were frozen using dry ice, with both the SOMOS 6110 and acetal molds first cooled to in a refrigerator freezer to reduce the potential for mold damage due to thermal shock [12].After 30 minutes in the dry ice, the molds were returned to room temperature (-20° C) and allowed to warm for approximately 4 minutes.

The molds are inverted and taped gently until the152still-frozen powder mass is released from the mold. The frozen powder mass is placed in a furnace boat on a layer of FPM tool powder and then subsequently surrounded with more tool powder before slurry melting begins.The frozen powder slurry is allowed to m~lt, as the furnace boat remains at ro~m temp~rature throughout storage and transport to the si~tenng furnace. The FPM tool powder IS suffICI~~t to retain tool powder shape dur~ng carri~r meltIn~. T~e presence. of the tool powder may also faCIlItate carrier removal through capIllary actIon by wlckIng the carner away from the part powder as has been demonstrated for debinding powder injection molded components [2].

Since sublimation is not required for carrier removal and since the part does not need to develop green strength to be self-supporting, no organic binder constituents are used in the carrier mixture.2.4 Carrier Removal & SinteringThe carrier used for freeze molding must be evaporated and the powder used for rapid tooling must be sintered. This is accomplished using a sintering furnace under protective atmosphere. Tooling was produced in both 316 stainless steel and nickel. For both powders, the sintering process was begun by heating the furnace to 85° C. This temperature was held for 1.5 hours. Then, the furnace temperature was increased to 115° C. This temperature was held for 1.5 hours to remove the carrier. Lengthy debinding times required for freeze drying sublimation are cited as a potential disadvantage of freeze molding [2]; however, it is unclear if the three hours used here is comparable to these times. Furnace temperature was then increased at 10° C/min.

Furnace temperature was then increased at 10° C/min. to a maximum sintering temperature of 1200° C for both 316 stainless steel and nickel powders. The powder was held at these temperatures for 45 minutes. The furnace cool-down commences at a rate no faster than 5° C/min. Sintering time is selected to minimize grain growth and is consistent with that recommended in the literature [13]. The nickel powder is sintered in an inert atmosphere of argon. The 316 stainless steel powder is sintered in a reducing atmosphere of dry hydrogen. The sintering cycles and atmospheres have not been optimized for the specific powders used.3. EXPERIMENTAL RESULTS & DISCUSSIONThe procedure described above was used to sinter dies for the example application of sheet metal forming. Fully dense material was not required for this demonstration; however, infiltration may be used in applications where improved mechanical properties or non-permeable tooling is required.

Fully dense material was not required for this demonstration; however, infiltration may be used in applications where improved mechanical properties or non-permeable tooling is required. This procedure should be suitable for producing prototype and short-run production tooling for a variety of other manufacturing processes such as injection molding, composite lay-up, sink-type EDM, extrusion, and thermoforming. The results presented below illustrate the potential utility of the FPM rapid tooling technique described here.3.1 Part EjectionPart ejection from the NC machined molds was relatively straight-forward. After allowing a short period for mold warming following freeze molding, gentle tapping is used to eject the resulting part. The part surface appeared quite smooth and even fine features on the molds, such as the marks from cutter rotation along tool paths, were visible on the molded part.Part ejection from the SLA molds was slightly more difficult.

Despite the use of mold release, a significantly longer duration was required to allow the molds to warm before parts could be ejected. While this may be due to different thermal conductivity of the mold materials or different geometries of the molds, inspection of the mold suggests that many of the 0.006" layers comprising the SLA part create undercuts which hinder part ejection. Figure 5 shows a magnified view of the near vertical mold walls. Despite the presence of a 3° draft angle on these walls, it is thought that significant thermal expansion of the frozen mold must occur before the part can be ejected.Upon ejection, many small fragments of the molded powder slurry (still frozen) are found resting loose on the top surface of the mold. This suggests that even thermal expansion of the mold does not fully release the frozen mass in all locations.

This suggests that even thermal expansion of the mold does not fully release the frozen mass in all locations. These fragments can be removed by suction153before they melt and their breakage from the molded part does not appear to significantly impact tooling performance for the example application.SLAMold Material0.1 emFigure 5 - SLA Mold Wall UndercutsSimilar artifacts likely exist with other layer-based SFF processes, and this illustration is not intended to suggest that the SLA process is unsuitable for this application. In fact, the SOMaS 6110 epoxy resin molds withstood both the thermal cycle of molding and impact required for part ejection without damage. In contrast to past experience with fragile parts made from early SLA resins, the SOMOSTM 6110 parts appeared to be quite tough.

In contrast to past experience with fragile parts made from early SLA resins, the SOMOSTM 6110 parts appeared to be quite tough. Constructing parts from finer resin layers or employing some additional finishing operation(s) may improve part ejection.3.2 Sintered DiesAfter sintering as described in section 2.4, the dies must be removed from the FPM tool powder in which they were embedded. This is accomplished simply by brushing the loose tool powder from the dies. Figure 6 shows the resulting nickel die set (left) produced by the SLA molds and the 316 stainless steel die set (right) produced by the NC machined molds.Figure 6 - Sintered Nickel Die Set from SLA Molds (left), and Sintered SS316 Die Set From NC Machined Molds (right)The layer-wise stair-stepping of the SLA molds transfers to the die-set, with some evidence of sharp corners of the stair-steps being damaged during mold ejection.

Although not clearly visible in Figure 6, the NC machined dies pick-up fine detail from the mold surface showing the concentric paths of the ball mill used to fabricate the mold protrusion. 3.2.1 Dimensional ChangeMaterial shrinkage resulting during consolidation caused dimensional reduction of both die sets. On average, the nickel shrunk 15.4% when measured linearly. The 316 stainless steel shrunk 3.40/0 when measured linearly. When material shrinkage is uniform it is possible to account for it simply by model scaling. Shrinkage experienced at prescribed processing conditions would typically be determined prior to mold design, and this shrinkage taken into account to over-size the mold geometries by an appropriate dimension.

In this experiment, the additional steps of shrinkage compensation were not performed and a smaller-than-designed mold geometry was accepted for demonstration purposes.Some non-uniform shrinkage was observed along the die axis held vertical during mold filling. This is believed to be caused by the development of density gradients in the compact as a result of mold filling. Deviations of 0.3% and 0.4% were observed on the female 316SS and nickel dies; however, the male dies exhibited deviations of 1.2% and 3% for the respective die material. This increased error is throught to result from poor mold filling due to the small mold gate154in the male die molds, as shown in Figures 3 and 4 above. Alternate gate geometries or improved mold filling techniques will likely eliminate this non-uniformity. 3.2.2 MicrostructureAnalysis of the microstructure co~firms the expected -. tooling.sintered from lo~se powder is rather porous.

3.2.2 MicrostructureAnalysis of the microstructure co~firms the expected -. tooling.sintered from lo~se powder is rather porous. Figure 7 shows the mIcrostructures of the smtered nIckel and 316 stamless steel die materials.."yl,," . ~" • r1.~ ~ ,.,....tc~~~~ " ~f}~"~' §t. .:L~*",~¥l~~;;. ~)~Z\l;i$'''~ F~;~.•~j;}.t:~~ .j; ..1..~el!,~~J :;,(~!~~~~!L~~~ .. -...""Figure 7 - Microstructure and Edge of (a)Nickel and (b)316 Stainless Steel FPM Rapid ToolingOptical density measurements using image processing techniques indicate densities of 80% of theoretical for the nickel dies and 60% of theoretical for the 316 stainless steel dies. However, density improvement should be possible with an optimized sintering schedule.These micrographs also illustrate the surface quality obtained on the near vertical edges (shown horizontally in this figure) of the tooling made by the SLA molds (a) and by the NC machined molds (b).

The rougher surface finish of the SLA molds transfers to the tooling. However, surface roughness as related to particle size must also be considered. The internal microstructure of the parts reflects the spherical morphology of the powders used to produce them. Note that both parts are porous and appear to be only stage one sintered. This enables application of infiltration or impregnation to realize fully-dense tooling, and either of these operations would serve to somewhat smooth the surface finish of the tooling. Pores resulting from gas entrapment during molding have been observed in some sintered material, and measures must be taken to eliminate these in practical situations.3.3 Metal FormingThe sintered dies were installed on a press and forming was attempted on 0.033" low-carbon mild sheet steel. This steel fractured during pressing on the unlubricated dies.

This steel fractured during pressing on the unlubricated dies. Examination of the metal flow of the sheet after forming suggests that high die friction near the fillet between the female die parting plane and the hemispherical depression is probably responsible for this forming failure. Instead of attempting to improve die surface finish manually, as is common even with production tooling, a 0.005" teflon film was used as an interpolator between both dies and the sheet metal during a second forming trial as suggested by [14]. This time the material did not fracture and forming was successful.

This time the material did not fracture and forming was successful. Figure 8 illustrates both the failed (a) and successful (b) forming trials using the SLA molded die set and the successful (c) forming trial using the NC machined molded die set.1 emFigure 8 - Metal Forming Results on 0.033" Sheet using (a) Unlubricated Dies from SLA Molds, (b) Dies from SLA Molds with Teflon Interpolator, and (c) SS316 Dies from NC Machined Molds155Although applying an interpolator sheet to the die surface instead of manual finishing may not be an economically viable alternative for high-volume part production, it is suitable for prototyping and low-volume work where minimizing die fabrication time is advantageous. Analyzing the affect of additional forming variables, such as pressing rate or binder tension were beyond the scope of this research effort.4. CONCLUSIONThe preceding results indicate that FPM/freeze molding is a useful secondary material conversion option for RP.

CONCLUSIONThe preceding results indicate that FPM/freeze molding is a useful secondary material conversion option for RP. This conversion route uses RP parts directly as tooling for powder shaping. This method can be used to quicly fabricate tooling from conventional materials as well as a number of high performance materials that can only be processed using powder metallurgy. This paper focused on sheet metal forming, however, a variety of other tooling such as injection molding cores and cavities, composite lay-up forms, EDM electrodes, extrusion dies, and thermoforming tooling may also be produced using this approach.Turn-around-time for the tooling featured in this report was relatively short.

In an ideal case,in-process time breaks down as follows:Day 1: Design formed part geometry and die geometry, fabricate RP molds that evening; • Day 2: RP molds in transit; • Day 3: Prepare powder slurry, fill & freeze molds, eject parts and embed in FPM tool powder, remove carrier and sinter over-night;Day 4: Remove solid tooling from embedding powder; use dies to press and form sample. Since mold filling and ejection are manual operations, and the resulting powder compact is fragile, more than one molding attempt may be required to realize a preform of suitable quality. If full density tools are required, additional time must be allocated for infiltrating, impregnating, or hot isostatic pressing.Additional investigation is certainly required to determine how the FPM/freeze molding approach described in this paper compares with alternative secondary material conversion options for RP and alternative rapid tooling methods.

It must also be contrast with direct methods for tooling production. The FPM rapid tooling procedure eliminates the need for the production of molds from RP patterns by directly creating the mold. This removes one shape conversion operation which reduces delay and may increase tooling precision. If layer-based SFF parts are used as molds, this technique still suffers the delay associated with layer-wise additive material build-up; however, this delay must only be incurred once during mold production. The resulting molds can be used to shape powder multiple times. Consequently, the "cost" of layer-wise fabrication may be amortized across multiple parts - whether these parts represent rapid tooling or end-use components. Alternatively, when relatively simple geometries are involved, NC machining can be performed quite efficiently on plastic mold material.

Alternatively, when relatively simple geometries are involved, NC machining can be performed quite efficiently on plastic mold material. Driven by appropriate CAM software, this is an RPIRT path which should not be overlooked.ACKNOWLEDGEMENTSThis research has been conducted at the New York State Center for Advanced Technology (CAT) in Automation, Robotics and Manufacturing at Rensselaer Polytechnic Institute and supported in part by Manufacturing Solutions, Inc. The CAT is partially funded by a block grant from the New York State Science and Technology Foundation.The authors wish to thank Gerd Beckmann, of Beckmann Engineering, for the initial suggestion of the use of freeze molding and Bill Carter, of General Electric's Corporate Research and Development Laboratory, for help with early freeze molding demonstrations.

Many of the results featured in this paper would have been impossible without the generous contribution of SOMOSTM 6110 epoxy SLA parts by Bronson Hokuf of DuPont's Somos Solid Imaging Materials Group. His generosity and helpful suggestions are greatly appreciated. Powders used in this work were kindly donated by Anval, Norton Industrial Ceramics Corp., and Novamet Specialty156Products. The authors are grateful for the assistance of Norm Gendron in preparing the metallography results featured in this paper.LITERATURE CITED [1]Rock, S. J. and C. R. Gilman, "A New SFF. Pr~~ess for. Functional Part ~aI2id Prototyping and Manufacturing: Freeform Powder Moldmg, Symposium Proceedings, H. L. Marcus, et. al. (eds.), pp. 80-87, The University of Texas at Austin, Austin, TX, 1995. German, R. M., Powder Injection Molding, Metal Powder Industries Federation, Princeton, New Jersey, 1990. Fussell, P. S. and L. E.

eeze Extrusion Fabrication of 13-93 Bioactive Glass Scaffolds for Bone RepairT.S. Huang,1* N.D. Doiphode,2 M.N. Rahaman,1 M.C. Leu,2 B.S. Bal,3 D.E. Day11Department of Materials Science and Engineering, and Center for Bone and Tissue Repair andRegeneration, Missouri University of Science and Technology, Rolla, Missouri 65409 2Department of Mechanical and Aerospace Engineering, Missouri University of Science andTechnology, Rolla, Missouri 654093Department of Orthopaedic Surgery, University of Missouri–Columbia, Columbia, Missouri,65212Reviewed, accepted September 23, 2010AbstractThere is an increasing demand for synthetic scaffolds with the requisite biocompatibility,internal architecture, and mechanical properties for the bone repair and regeneration.

In this work,scaffolds of a silicate bioactive glass (13-93) were prepared by a freeze extrusion fabrication(FEF) method and evaluated in vitro for potential applications in bone repair and regeneration.The process parameters for FEF production of scaffolds with the requisite microstructuralcharacteristics, as well as the mechanical and cell culture response of the scaffolds wereevaluated. After binder burnout and sintering (60 min at 700°C), the scaffolds consisted of adense glass network with interpenetrating pores (porosity ≈ 50%; pore width = 100−500 µm).These scaffolds had a compressive strength of 140 ± 70 MPa, which is comparable to thestrength of human cortical bone and far higher than the strengths of bioactive glass and ceramicscaffolds prepared by more conventional methods. The scaffolds also supported the proliferationof osteogenic MLO-A5 cells, indicating their biocompatibility.

The scaffolds also supported the proliferationof osteogenic MLO-A5 cells, indicating their biocompatibility. Potential application of thesescaffolds in the repair and regeneration of load-bearing bones, such as segmental defects in longbones, is discussed.Keywords: Freeze extrusion fabrication; bioactive glass; solid freeform fabrication; bone repair *Corresponding author. E-mail: hts@mst.edu (T.S. Huang)9321. IntroductionThere is an increasing demand for synthetic scaffolds with the requisite biocompatibility,mechanical properties and internal architecture for the repair and regeneration of tissues andorgans.

An ideal scaffold for bone repair and regeneration should have the followingcharacteristics [1]: (1) biocompatible (non-toxic) with the ability to promote cell adhesion,proliferation, and differentiation; (2) porous three-dimensional (3D) architecture to allow cellproliferation, vascularization, and diffusion of nutrients between the cells seeded within thematrix and the surroundings; (3) mechanical properties comparable to the bone to be replaced; (4)ability to bond firmly to bone and soft tissue; (5) degradation rate similar to the rate at whichnew bone is formed; (6) processability into the desired anatomical shape.Bioactive glass has several attractive properties for application as a scaffold material[2,3]. Bioactive glass reacts with the body fluids, forming a surface layer of hydroxyapatite (themain mineral constituent of bone) which is responsible for forming a firm bond with hard andsoft tissues.

Bioactive glass is osteoconductive as well as osteoinductive, and has a widelyrecognized ability to support new bone growth. However, porous bioactive glass scaffoldsfabricated using conventional methods commonly have low strength (<20 MPa) [3,4], so theiruse in the repair and regeneration of load-bearing bones is challenging. Attempts are currentlybeing made to improve the mechanical properties of bioactive glass scaffolds for the repair ofload-bearing bones [4-6]. These attempts are focused mainly on the control of the pore shape,pore size, and pore orientation to improve the strength of the scaffolds.Solid freeform fabrication (SFF) can potentially produce scaffolds with customizedexternal shape, as well as predefined and reproducible internal architecture (porosity; pore size;933and pore distribution).

An advantage of the SFF technology is the potential ease of creatingscaffolds with the anatomical shape and dimensions tailored specifically to individual patients.SFF technologies involve building three-dimensional (3D) objects layer by layer. Although thereare several variants of SFF technology, the general process involves producing a computer-generated model, which represents the physical prototype to be built, using computer-aideddesign (CAD) software. The model is next converted into a format that is analyzed by acomputer, which slices the model into cross-sectional layers. The data is then implemented to theSFF machine which systematically produces the physical prototype layer by layer.

The data is then implemented to theSFF machine which systematically produces the physical prototype layer by layer. Post-processing of the formed article is often necessary to remove temporary support structures orprocessing aids, particularly in the case of ceramic and glass particles that are commonlydifficult to bond directly by most available SFF techniques.The most widely researched SFF techniques include fused deposition modeling (FDM) ofpolymeric materials [7,8], fused deposition of ceramics (FDC) [9-11], selective laser sintering(SLS) [12,13], 3-D printing [14], stereolithography [15], and robocasting [16]. Freeze extrusionfabrication (FEF) has received less attention, but the technique has been used to fabricate denseceramic articles for mechanical engineering applications [17]. FEF is based on a combination oftechniques taken from fused deposition, freeze casting, and robocasting [18].

FEF is based on a combination oftechniques taken from fused deposition, freeze casting, and robocasting [18]. In FEF, an aqueousmixture with a paste-like consistency is extruded through and orifice to form filaments, whichare frozen in situ to avoid slumping or distortion of the formed article. The rheology of the pasteis critical for providing the plastic properties required for avoiding flaws in the formed article.After formation by FEF, the article is subjected to a freeze-drying step to sublime the frozenliquid.934In FEF of ceramic or glass, particles are commonly mixed with a polymeric binder phaseto provide the required plastic properties for extrusion. FEF of these materials involve key post-processing steps. Following freeze drying to remove the frozen liquid, a binder removal step isnecessary prior to a sintering step in which the particles are thermally bonded to form a strongnetwork.

Following freeze drying to remove the frozen liquid, a binder removal step isnecessary prior to a sintering step in which the particles are thermally bonded to form a strongnetwork. Ideally, the binder phase used to impart the plasticity to the extrudate must be removedcompletely (typically by thermal decomposition) prior to the onset of the sintering step to avoidthe presence of impurity phases in the final article. In the sintering step, the article is heated to asufficiently high temperature to cause matter transport in order to bond the particles into a densestrong network. In the case of ceramics, the sintering temperature is typically 0.5−0.9 of themelting temperature, whereas the sintering temperature is typically between the glass transitiontemperature and the melting temperature for glass.In the last decade, SFF methods have been widely applied to the production implants andscaffolds for biomedical applications.

The biomaterials used in these studies have been mainlybiodegradable polymers such as poly(glycolic acid), poly(lactic acid) and their copolymers, bio-inert metals such as titanium, and bioactive ceramics such as hydroxyapatite [19-28]. There isgrowing interest in the production of bioactive ceramics and bioactive glass scaffolds forpotential bone repair applications because their attractive properties.The objective of this work was to evaluate the feasibility of forming bioactive glassscaffolds using an FEF technique. The development of extrudable paste with optimized rheologyand the process parameters for the production of scaffolds with pre-determined internalarchitecture were studied. A silicate bioactive glass designated 13-93 was used in this workbecause of our previous experience with forming scaffolds of this glass using more conventional935techniques.

A silicate bioactive glass designated 13-93 was used in this workbecause of our previous experience with forming scaffolds of this glass using more conventional935techniques. Products of 13-93 bioactive glass are also approved for in vivo use in the UnitedStates, Europe, and elsewhere [29].2. Experimental Procedures2.1. Preparation of starting materials and extrudable pasteThe bioactive glass used in this work is designated 13-93, with the composition (wt%):53.0 SiO2, 6.0 Na2O; 12.0 K2O; 6.0 MgO, 20.0 CaO; 4.0 P2O5. Coarse particles of the glass werekindly provided by Mo-Sci Corp., Rolla, Missouri. The as-received glass particles were groundto a size finer than ∼3 µm, using a combination of two methods.

Coarse particles of the glass werekindly provided by Mo-Sci Corp., Rolla, Missouri. The as-received glass particles were groundto a size finer than ∼3 µm, using a combination of two methods. Particles of size smaller than 45µm were first obtained by grinding the as-received particles in a steel shatterbox (SPEXSamplePrep LLC, Metuchen, NJ) and sieving the particles through a 325 mesh sieve.Approximately 100 g of these particles were then dispersed in deionizedwater and ground for 1 hin an attrition mill (Model 01-HD, Union Process, Akron, OH) using ZrO2 grinding media. Afterevaporation of the water from the attrition-milled slurry using a hot plate, the particles were driedfor 24 h in oven at 65°C. The dried, agglomerated powder was ground in an agate mortar andpestle, and sieved through a 325 mesh sieve.An extrudable paste for use in the FEF process was prepared as follows.

The dried, agglomerated powder was ground in an agate mortar andpestle, and sieved through a 325 mesh sieve.An extrudable paste for use in the FEF process was prepared as follows. First, anaqueous-based slurry with the composition given in Table I, consisting of glass particles,deionized water, and polymeric additives, was formed into a homogeneous mixture by ballmilling for 24 h using Al2O3 milling media. These polymeric additives were selected becausethey were used previously in the fabrication of Al2O3 articles by the FEF method [18]. Second,the ball-milled slurry was heated under a controlled temperature/time schedule (70 min at 65°C)to evaporate some of the water, and to control the viscosity of the paste. Pastes with different936viscosities were tested in the FEF equipment to determine whether they were extrudable, and thesystem with the optimum viscosity for extrusion was chosen by a trial-and-error method.Table I.

Composition of starting slurry used in the preparation of the extrudable paste for FEF.ComponentConcentration (vol%)FunctionManufacturer13-93 glass particles40.00Solid phase Mo-Sci Corp., Rolla, MOEasySperse0.50DispersantISP Technologies, Inc., Wayne, NJSurfnolGlycerol0.501.00Defoamer WCCA*Air Products & Chemicals, Inc., Allentown, PAAlfa Aesar, Ward Hill, MAPEG 4001.00LubricantAlfa Aesar, Ward Hill, MAAquazol 54.00BinderISP Technologies, Inc., Wayne, NJDeionized water *WCCA: Water crystallization control agent53.00Solvent—2.2 Freeze extrusion fabrication of bioactive glass scaffoldFigure 1 shows images of the main components of the FEF equipment used in this work.The paste to be extruded is loaded into a heated reservoir housed in a computer-controlled gantrythat allows programmable movement in three dimensions.

A nozzle attached to the reservoirdetermines the diameter of the extrudate which is deposited on a fixed platform cooled to therequired temperature using liquid nitrogen. The system is contained within a cooled chamber.In the present work, the fabrication environment temperature was −20°C. The paste wasextruded layer by layer, with each layer deposited at 90° relative to the preceding layer. Based onthe rheology of the paste, the software in the FEF machine was adjusted to control the extrusionforce and rate of deposition. A nozzle diameter of 580 µm was used in these experiments, andthe spacing between adjacent filaments was 600 µm.937Figure 1. Images showing the main components of the freeze extrusion fabrication equipment used in the present work.2.3. Post processing of as-formed FEF constructThe decomposition kinetics of the polymeric binder (without glass particles) werefollowed using thermogravimetric analysis (STA409, NETZSCH).

Post processing of as-formed FEF constructThe decomposition kinetics of the polymeric binder (without glass particles) werefollowed using thermogravimetric analysis (STA409, NETZSCH). Based on these kinetic data, atemperature−time schedule was developed for the removal of the polymeric additives from theas-formed FEF construct. The objective was to completely remove the binder by thermaldecomposition prior to sintering of the glass particles. Incomplete removal of the binder prior tosintering resulted in scaffolds with a black color, due to residual carbon entrapped in the glass. Aschedule lasting 5 days, consisting of a slow heating rate (3−5°C/h) and several isothermalholding stages, resulted in complete removal of the binder. Following the binder removalschedule, the construct was sintered in air for 1 h at 700°C (heating rate = 5°C/min) to densifythe glass phase.9382.4.

Following the binder removalschedule, the construct was sintered in air for 1 h at 700°C (heating rate = 5°C/min) to densifythe glass phase.9382.4. Structural and mechanical evaluation of 13-93 bioactive glass scaffoldsFollowing the sintering step, bioactive glass scaffolds were ground to form a powder(<45 µm) and analyzed using X-ray diffraction, XRD (Model D/mas 2550 v; Rigaku, TheWoodlands, TX). The analysis was performed using Cu Kα radiation (λ = 0.15406 nm) at ascanning rate of 0.01° 2θ/min in the range 3−90° 2θ. The microstructure of the scaffold and theglass phase was examined using conventional methods in an optical microscope and a scanningelectron microscope, SEM (S-4700; Hitachi, Japan).The mechanical response of the fabricated scaffolds in compression was measured at acrosshead speed of 0.2 mm/min in an Instron machine (Model 4205; Instron, Norwood, MA).

Sixcube-shaped samples (5 mm in length), surface ground using a diamond coated wheel andsectioned using a diamond-coated blade from the fabricated scaffolds, were tested. Thecompression force was applied along the thickness direction of the scaffold.2.5. Cell cultureThe ability of the fabricated scaffolds to support the proliferation of MLO-A5 cells, anestablished murine osteogenic cell line, was used to confirm their biocompatibility. The MLO-A5 cells were kindly provided by Professor Lynda F. Bonewald, University of Missouri-KansasCity. After they were washed twice with water and dried, cube-shaped scaffolds (5 mm inlength), similar to those used the mechanical tests described above, were sterilized by heating for2 h at 500°C, seeded with 50,000 MLO-A5 cells suspended in 100 µl of complete medium andincubated for 4 h to permit cell attachment. The cell-seeded scaffolds were then transferred to a24-well plate containing 2 ml of complete medium per well.

The cell-seeded scaffolds were then transferred to a24-well plate containing 2 ml of complete medium per well. All cell cultures were maintained at37 ºC in a humidified atmosphere of 5% CO2, with the medium changed every 2 days.939To visualize the metabolically active cells on and within the scaffolds, each cell-seededscaffold was placed in 0.2 ml serum-free medium containing 0.1 mg of the tetrazolium salt MTTfor the last 4 h of incubation. After incubation, the scaffolds were briefly rinsed in PBS, blotted,and allowed to dry. Images of the scaffolds were obtained using a stereomicroscope fitted with adigital camera to qualitatively assess the distribution of insoluble purple formazan, a product ofmitochondrial reduction of MTT by viable cells.3. Results and Discussion3.1.

Results and Discussion3.1. Structural characteristics of bioactive glass scaffoldsFigure 2 shows optical images of 13-93 bioactive glass constructs, as-formed by the FEFtechnique (Figure 2a), and after sintering at 700°C (Figure 2b). The as–formed construct(approximately 36 mm × 30 mm × 6 mm thick) was obtained by extrusion of the paste thoroughan orifice of diameter 580 µm. The diameter of the deposited filaments (∼600 µm) was slightlylarger than the diameter of the orifice due to expansion of the plastic paste upon extrusion. Asshown, the width of the pores in the plane of the deposition was ∼500 µm and the porosity of theas-formed FEF construct was ~75% (including the porosity of the glass filaments).There was no measurable shrinkage of the FEF scaffold during the freeze drying step,and the shrinkage during the binder burnout step was small (linear shrinkage <3%).

However, theconstruct shrank considerably during the sintering step, as a result of the densification of thenetwork of fine glass particles. The linear shrinkage was ~28% in the plane of deposition, and∼20% in the thickness direction of the scaffold. The sintered scaffolds consisted of glassfilaments with a thickness of ∼300 µm, porosity of ~50%, and pore width (in the plane of thedeposition) of 300 µm (Table II).940Figure 2. Optical images showing a bioactive glass construct (a) as-formed by the FEF technique, and (b) after binder burnout and sintering.Table II.

Optical images showing a bioactive glass construct (a) as-formed by the FEF technique, and (b) after binder burnout and sintering.Table II. Structural characteristics of bioactive glass scaffolds as-formed by FEF and after sintering at 700°C.ParameterFEF scaffoldSintered scaffoldFilament diameter (µm)Porosity (%)600 75*30050Pore width (µm)500300Porosity including the porosity in the filaments before sinteringAs Figure 2 shows, the FEF construct retained its shape and architecture during the post-processing steps (freeze drying, binder burnout; sintering). SEM images of fractured crosssections (Figure 3) showed that sintering resulted in almost complete densification of the glassphase, and the glass network contained only a few fine, isolated pores. Furthermore, there wasgood bonding between the adjacent layers of the sintered scaffold. These structuralcharacteristics are desirable for enhancing the overall strength of the sintered scaffold.

Furthermore, there wasgood bonding between the adjacent layers of the sintered scaffold. These structuralcharacteristics are desirable for enhancing the overall strength of the sintered scaffold. Theability to achieve nearly full density presumably resulted from the fine particle size of the941starting glass particles (<3 µm), the homogeneous mixing of the particles and binder phase in thepaste used in the extrusion, and high packing density of the glass particles in the paste.Figure 3a shows that the pore width (100−150 µm) in the thickness direction of thesintered scaffold was smaller than that in deposition plane (~300 µm). This smaller pore width inthe thickness direction resulted presumably from deformation due to gravitational effects duringthe FEF step and particularly during the sintering step.

This smaller pore width inthe thickness direction resulted presumably from deformation due to gravitational effects duringthe FEF step and particularly during the sintering step. In sintering, the reduction in the viscosityof the glass, necessary to achieve viscous flow densification of the glass phase and bondingbetween adjacent layers of the scaffold, also enhanced deformation of the glass by creep underthe force of gravity.Figure 3. SEM images of the fractured cross section of a sintered bioactive glass scaffold showing (a) good bonding between adjacent layers of the scaffold, and (b) almost full densification of the glass phase.

(The section was along the thickness direction of the scaffold.)X-ray analysis of the sintered scaffolds did not show any diffraction peaks (Figure 4).Instead, the XRD pattern showed a broad band (centered at ∼30° 2θ) characteristic of a glass.The ability to avoid crystallization of the bioactive glass during the scaffold fabrication processis beneficial for achieving high density of the glass network in the scaffold (and, hence forachieving high strength) and for maintaining the bioactive potential of the glass. Crystallizationduring sintering leads to a glass−ceramic material which is often difficult to densify. Furthermore,942while the glass−ceramic still retains the ability to form a hydroxyapatite surface layer, the rate offormation of the hydroxyapatite layer is markedly reduced.Figure 4. X-ray diffraction pattern of sintered bioactive glass scaffold, showing that the glass remained amorphous after the sintering step.3.2.

X-ray diffraction pattern of sintered bioactive glass scaffold, showing that the glass remained amorphous after the sintering step.3.2. Mechanical response of sintered bioactive glass scaffoldsFigure 5a shows cube-shaped samples that were used in tests to measure the mechanicalresponse of the sintered scaffolds in compression. The scaffolds showed a brittle response,typical of dense ceramics and glass (Figure 5b). The applied compressive stress increased almostlinearly with the deformation until the sample failed in a catastrophic manner into several pieces.For the 6 samples tested, the compressive strength of the scaffold was 140 ± 70 MPa. Thisaverage strength is in the range of values reported for human cortical bone (120−180 MPa).

Thisaverage strength is in the range of values reported for human cortical bone (120−180 MPa). Thelarge standard deviation in the compressive strength may result from sample non-uniformity dueto the small size of the samples tested and from macroscopic flaws arising from the FEF process.Since the linear dimension of the samples was 5 mm, variable in the sample uniformity can havea marked effect on the mechanical properties. The elastic modulus of the samples, determinedfrom the slope of the stress vs. strain data, was 5−6 GPa, which is lower than the elastic modulus943of cortical bone (10−20 GPa).Figure 5. (a) Cube-shaped samples (5 mm in length) used in mechanical testing; (b) mechanical response (applied stress vs. deformation) of a sample in compression.Table III shows a comparison of the compressive strengths of biodegradable polymer,bioactive glass, and hydroxyapatite scaffolds prepared by various methods [30].

The valuesshown in Table III are not meant to be exhaustive. Instead, they show the results of selectedstudies for various biomaterials and scaffold fabrication techniques. Although the compressive944strength in the present study showed a large standard deviation, the lowest strength (60 MPa) isstill far higher than the strengths of biodegradable polymer scaffolds, as well as bioactive glassand hydroxyapatite scaffolds prepared by more conventional methods.Table III.

The purplepigment visible on the scaffold is an indication of viable cells, and is the result of mitochondrialreduction of MTT to an insoluble formazan product. The increase in the intensity of the purplecolor with time indicated the proliferation of viable, metabolically active cells on the scaffolds.945Figure 6. Optical images showing cell-seeded bioactive glass scaffolds treated with MTT after culture intervals of (a) 2 days; (b) 4 days, and (c) 6 days. The increase in intensity of the purple color (arrow) indicates the ability of the scaffolds to support cell proliferation.The present work shows promising results for using the FEF technique in the fabricationof bioactive glass scaffolds for the repair and regeneration of load bearing bones.

In addition tothe benefit of other SFF methods in producing articles with predesigned external shape andinternal architecture, the FEF was shown to produce porous 3D scaffolds with an almost fullydense glass network, which resulted in a compressive strength comparable to that of corticalbone.4. ConclusionsThe present results showed the feasibility of a freeze extrusion fabrication (FEF)technique for creating bioactive glass scaffolds with the requisite pore architecture andmechanical strength for potential application in the repair and regeneration of load-bearing bones.An extrudable paste was developed for the production of scaffolds with uniform shape andinternal architecture. The scaffolds retained their external shape and internal architecture duringthe FEF process, as well as during the post-processing steps (freeze drying; binder removal;sintering).

The scaffolds retained their external shape and internal architecture duringthe FEF process, as well as during the post-processing steps (freeze drying; binder removal;sintering). Sintered scaffolds of 13-93 glass with a porosity of ∼50%, and interconnect pores ofsize 300 × 300 × 100−150 µm, had a compressive strength of 140 ± 70 MPa, comparable to the946strength of human cortical bone. These sintered scaffolds showed a brittle mechanical responsein which the stress increased linearly with deformation until catastrophic failure. The scaffoldsremained amorphous after sintering, and supported the proliferation of osteogenic MLO-A5 cells,showing their biocompatibility.References1. D.W. Hutmacher, “Scaffold design and fabrication technologies for engineering tissues –state of the art and future perspectives,” J. Biomater. Sci. Polymer Edn, 12 [1] 107-12 (2001).2. L.L. Hench, “Bioceramics,” J. Am. Ceram. Soc. 81 1705-28 (1998).3. M.N. Rahaman, A. Yao, B.S.

eeze-form Extrusion Fabrication of CeramicsTieshu Huang, Michael S. Mason, Gregory E. Hilmas, and Ming C. Leu1870 Miner CircleDepartment of Materials Science & EngineeringUniversity of Missouri – Rolla, Missouri 65409–0050Email: {hts,mmason,ghilmas,mleu}@umr.eduAbstractA novel, environmentally friendly solid freeform fabrication method called Freeze-form Extrusion Fabrication (FEF) has been developed for the fabrication of ceramic-based components. The method is based on deposition of ceramic pastes using water as the media. The ceramic solids loading can be 50 vol. % or higher and initial studies have focused on the use of aluminum oxide (Al2O3). The FEF system components and their interaction are examined, and the main process parameters affecting part geometry defined. 3-D shaped components have been fabricated by extrusion deposition of the ceramic paste in a layer-by-layer fashion.

3-D shaped components have been fabricated by extrusion deposition of the ceramic paste in a layer-by-layer fashion. The feasibility of this process has been demonstrated by building components having a simple geometry, such as cylinders and solid or hollow cones. Hollow cones have also been fabricated to demonstrate the ability to build structures with sloped walls.I. IntroductionAdvanced ceramics are needed for applications in the aerospace, automotive, and other industries [1]. Certain ceramic materials can meet the high temperature requirements, but processing these materials into usable components is often challenging, costly, and time consuming. Being able to build a ceramic part directly by the process of material addition could considerably reduce the cost and fabrication time. Many Solid Freeform Fabrication (SFF) processes are capable of producing ceramic parts directly.

Many Solid Freeform Fabrication (SFF) processes are capable of producing ceramic parts directly. One of the most extensively researched methods for rapid prototyping of ceramics is Fused Deposition of Ceramics (FDC) [2-5]. Fused Deposition Modeling (FDM) is a SFF technology developed by Stratasys (Eden Prairie, Minnesota) [5]. FDM is a layered manufacturing technique that extrudes a thin bead of plastic, one layer at a time. A thread of plastic is fed into an extrusion head, where it is heated into a semi-liquid state and extruded through a nozzle onto the previous layer of material. FDC is modified from the FDM technique by changing the deposition material from thermoplastics to a ceramic-thermoplastic mixture. FDC uses a relatively large amount of organic chemicals as binders (40-50%), providing the ability to process at low temperatures (< 200oC) and pressures (<500 psi), while solidifying rapidly to form a rigid backbone after extrusion.

However, removal of organic binders produces toxic and non-decomposable wastes, which are undesirable.72In FDM, The extrusion occurs at a plastic softening temperature to form an extrudable material. Some other extrusion deposition techniques have also been investigated, for example, aqueous porcelain paste extrusion to form artificial human teeth [7] and extrusion freeform fabrication of functional graded materials [8]. In the current study, Freeze-form Extrusion Fabrication (FEF) is being combined with a previously developed Rapid Freeze Prototyping (RFP) process. RFP builds a three-dimensional ice part according to its CAD model by depositing and rapidly freezing water droplets in a layer-by-layer manner. Compared with other SFF processes, it has many advantages including less expensive equipment and material, a cleaner material and process, less energy consumption, and a better surface finish. The process for fabricating ice parts is conducted in a freezer.

The process for fabricating ice parts is conducted in a freezer. After being ejected from the drop-on-demand nozzle, water droplets deposit on a substrate or previously formed layer of ice. The water droplets do not solidify immediately but unite to become part of the continuous water line, which is called the water deposit. Meanwhile, the water deposit changes phase to ice in a continuous fashion [9-12].Similar to FDC, the FEF process is also designed to use ceramic pastes with a high solids loading (≥ 50 vol.%). However, the pastes use only trace amounts of organic binder (1-4%). This is possible by performing the deposition within a freezer. The lower temperatures allow the slurry to maintain its shape by freezing the water present in the slurry.This paper will introduce the FEF process, the hardware that is currently being used, and the process parameters that affect the material deposition and the motion algorithm utilized to build hollow cone geometry.

Paste behavior, freeze drying, binder removal, sintering, and materials’ properties will also be detailed.II. Experimental Procedure1. FEF equipment setupThe FEF setup is composed of five major components: the motion system, control system, deposition device, force sensor, and freezer. The following sections give a brief overview of the equipment that makes up the FEF system. Figure 1 shows the FEF system outside of the freezer.1.1. Motion systemThree extended temperature range (down to - 40oC) stepper motors (Empire Magnetics, Rohnert Park, California) are used in conjunction with linear axes. Each linear axis has 254 mm of travel. The stepper motors have a stepping angle of 1.8 degrees, but have the ability to utilize “micro steps.” The motors connected with a lead screw provide a maximum velocity of 250 mm/s and can vary their micro steps from 1-250 giving the highest resolution of 0.0072 degrees/step.

In addition, the motors each have a resolver that feeds a signal to a Resolver-to- Digital converter (RDE), which converts the resolver signal into an equivalent encoder feedback, allowing for an axis resolution of 0.5 µm per step.73Figure 1: Freeze-form Extrusion Fabrication equipment setup1.2. Real-time control systemThe hardware portion of the control system is composed of a Pentium IV 3.03 GHz computer for software development and graphical user interface, which is programmed using LabView software (National Instruments, Austin, Texas). The chassis is a National Instruments PXI/SCXI 1010. The chassis has 8 slots for PXI (data acquisition and control) cards and 4 slots for SCXI (signal conditioning) cards. The cards are connected in the chassis backplane allowing for synchronous timing.

The cards are connected in the chassis backplane allowing for synchronous timing. The National Instruments PXI 8187 real-time controller has a dedicated 2.2 GHz processor with 1 GB of RAM and is equipped with a 30 GB hard drive, one General Purpose Interface Bus (GPIB), and two serial port connections. The controller was upgraded to an extended temperature range design due to excess heat created by the PXI cards.A National Instruments PXI-7334 motion control card is used to regulate the motors and read the encoder feedback signals. The card can control up to 4 individual axes operated by either servo or stepper motors. The maximum pulse rate for the stepper motors is 4 MHz and the maximum encoder feedback rate is 20 MHz for each axis. There are three limits: home, forward, and reverse. The card has a trajectory loop update rate of 62.5 µs.

There are three limits: home, forward, and reverse. The card has a trajectory loop update rate of 62.5 µs. There are also four single ended 8 bit TTL ports for auxiliary digital communication.A National Instruments PXI-6025E multifunction data acquisition card is used for data input and output. It has two 24 bit counter/timers, two 12 bit analog outputs, 32 digital I/O lines, and 16 analog input lines which can be used as 16 single ended or 8 double ended inputs. The maximum sampling rate is 200 kS/s per channel with the input operation range from a minimum range of -0.05 to 0.05 V up to a maximum range of -10 to 10 V.741.3. Deposition deviceAn extrusion device (Figure 2) is necessary for deposition of ceramic pastes having a high solids loading, as the actuator can apply the necessary amount of pressure.

Deposition deviceAn extrusion device (Figure 2) is necessary for deposition of ceramic pastes having a high solids loading, as the actuator can apply the necessary amount of pressure. The extrusion device is attached to the z-axis of the FEF system, and a stepper motor with a resolution of 16,000 steps/inch is used to actuate a linear axis for the extrusion process. The linear axis moves an extrusion ram, which applies pressure to a reservoir containing the paste, and ultimately to the extrusion nozzle, forcing material through the nozzle. The deposition rate can be changed by adjusting the stepper motor speed. A 60 cm3 extruder reservoir is used to hold and eject the ceramic paste. Nozzles of varying diameter ranging from 190 to 580 µm are attached to the end of the extruder to allow for different diameters and thus thicknesses of material extrudate.

Nozzles of varying diameter ranging from 190 to 580 µm are attached to the end of the extruder to allow for different diameters and thus thicknesses of material extrudate. .Stepper Stepper Motor MotorLinear Linear Axis AxisLoad cellExtrusion RamNozzleFigure 2: Extrusion device as connected to the z-axis.1.4. Force sensorA LC301-1k load cell from Omega Engineering (Stamford, Connecticut) has been implemented into the system. The load cell has been attached at the end of the extrusion linear axis. The load cell was chosen for force feedback that will allow for constant material deposition. The load cell measures the applied force of the linear axis on the extrusion ram. A differential voltage is sent by the load cell and read by the data acquisition board that corresponds to a force between 0 – 1000 lbs.2. Materials and ceramic paste preparationThe materials used to form the alumina (Al2O3) based ceramic pastes in the current study are listed in Table 1.

Materials and ceramic paste preparationThe materials used to form the alumina (Al2O3) based ceramic pastes in the current study are listed in Table 1. The solids loading of the pastes were from 45 ∼ 55 vol.%. The solids- loading affects the rheological properties of the paste, which directly influence the extrusion behavior. Aquazol is a water soluble binder with a pH of 7, which will not affect the paste’s pH75value. Binder additions of 1 ∼ 4 vol.% were added to the paste to achieve consistent extrusion behavior and to assist in forming a stronger green body after drying. PEG was added as an extrusion lubricant to reduce friction between the paste and the nozzle wall as well as the friction between the particles during extrusion.

PEG was added as an extrusion lubricant to reduce friction between the paste and the nozzle wall as well as the friction between the particles during extrusion. Glycerol was used to prevent the growth of large ice crystals and freezing defects associated with water crystallization [13], Davan C was used as a dispersant and Deionized water as the medium.Table 1: Materials used in paste preparationNo.Name and GradePropertiesManufacturer1622 (A-16SG) Al2O3APS 0.4 µmAlcoa2Poly(ethylene glycol), PEG-400MW 400Aldrich3Darvan CR. T. Vanderbilt4Aquazol 50MW 50,000ISP Technologies, Inc.5GlycerolAldrich6Deionized WaterThe Al2O3 powder, PEG, glycerol, Darvan C, and water are mixed in a Nalgene bottle and ball milled for 24 hours using alumina milling media. Ball milling is used to break up any agglomerates in the powder to produce a uniform mixture.

Ball milling is used to break up any agglomerates in the powder to produce a uniform mixture. Aquazol is dissolved in water at 60 °C using magnetic stirring to form a 50 vol.% Aquazol solution. After ball milling the Aquazol solution is added into the batch mixture using a vacuum mixer (Whip Mix, Model F, Louisville, KY) to form a uniform paste that does not contain bubbles. The final viscosity was adjusted by adding acid to control the pH of the paste. The resulting paste is then collected into syringes and is ready for FEF processing.3. Material DepositionPath planning and material deposition are affected by three main parameters: nozzle diameter, XY table speed, and extrusion rate. Currently three different nozzle diameters have been used in this study (190, 250, 580 µm). There is a tradeoff between part build time and part accuracy that must be chosen, either a fast build time with slightly less accurate part dimensions or high dimensional accuracy with a longer build time.

The extrusion rate and XY table speed must be matched in order to deposit material in a smooth continuous fashion. If the table speed is too high, the extruded material will have gaps in deposition. A high table speed creates a shearing force that overcomes the cohesion properties of the slurry. An example of this is shown in Figure 3. If the table speed is too slow then there is additional material buildup, which is undesirable for two reasons. The part accuracy is reduced do to a larger layer height and it may create the possibility of clogging the nozzle. If the nozzle is immersed in the deposited ceramic slurry for extended periods, the material starts to solidify within the nozzle thereby reducing the material flowrate and eventually causing complete clogging of the nozzle.76Figure 3: Example of gaps in material depositionThe extrusion rate is dependent upon the applied force from the extrusion ram. Force feedback is chosen to develop controllable material deposition.

Force feedback is chosen to develop controllable material deposition. A load cell has been attached to the end of the extrusion ram in order to measure the amount of force being applied to the extruder. This force is fed back to the PC. A basic on/off control algorithm has been written that keeps the applied force within +/- 10 lbs of the desired amount. A better control algorithm that will utilize the relationship between the extrusion velocity and applied force will be implemented later to improve material deposition.4. Freeze drying, binder removal, and sinteringFreeze drying was performed in a chamber that was set in a freezer and pumped with a mechanical vacuum pump. The freeze drying temperature was set at – 16 °C and the time is from 2 to 5 days in order to remove the water by sublimation, depending on the size of the sample.

The freeze drying temperature was set at – 16 °C and the time is from 2 to 5 days in order to remove the water by sublimation, depending on the size of the sample. Following freeze drying, the samples were dried further at ambient and elevated temperatures to complete the drying process.A thermogravimetric analyzer (Shimadzu TGA50, Lenexa, KS) was employed for determining the binder thermolysis behavior. Samples of the batched pastes were tested using a heating rate of 5 °C/min to 600 °C in air. A binder removal profile was then developed based on the thermogravimetric analysis (TGA) results and performed in a box furnace. The sintering temperature was 1550 °C, at a heating rate of 5 °C/min, and a hold time of 2 hours. Sample densities were obtained using the Archimedes method. The samples were sectioned in the direction shown in Figure 4 to check the sample uniformity, microstructure, and pore distribution using scanning electron microscopy (SEM) techniques (Jeol 330, Peabody, MA).

SEM micrographs were obtained from areas a and b, as shown in Figure 4.AabA-AFigure 4: Schematic drawing showing the sample sectioning position and SEM analysis areas.77III. Results and Discussion1. RheologyBecause Aquazol 50 is a neutral binder, it will not change the slurry/paste pH value when added; so the effects of dispersant content were tested prior to binder additions. The effect of binder content on viscosity of the paste was performed on batches containing 1.5 vol.% dispersant. Binder content was varied from 1 to 6 vol.% with a 1 vol.% resolution. The test results (Figure 5) show that the viscosity decreases as dispersant concentration increases when the content is less than 2 vol.%. However, the viscosity will increase as the dispersant content increases when the dispersant content is larger than 2 vol.%. It also can be determined that the slurry exhibits a strong shear thinning behavior for all ranges of dispersant content.

It also can be determined that the slurry exhibits a strong shear thinning behavior for all ranges of dispersant content. Dispersant content of paste recipe is decided by uniform ball milling and relatively high viscosity after binder additions. The relatively high viscosity is needed for extrusion. Figure 6 shows the effect of binder content on the paste viscosity. The tendency is that the viscosity increases as the binder content increases. This effect is significant in the low shear rate region, but not as significant in the high shear rate region. 2 to 4 vol.% binder was adapted in the paste preparation procedures since these pastes exhibited a low enough viscosity in the high shear rate range to be extrudable at low pressures, while having a high enough viscosity in the low shear rate range to quickly become rigid and provide green strength after extrusion. The pH of the paste is adjustable through addition of a 4% HNO3 water solution.

The pH of the paste is adjustable through addition of a 4% HNO3 water solution. Test results showed that pH adjustment was the most significant method for developing a uniform paste and for controlling paste viscosity.120) s .a P M 0 1 (310080601 vol.% 1.5 vol.% 2 vol.% 2.5 vol.% 3 vol.%y t i s o c s i V40200020406080100Shear rate (s-1)Figure 5: Effects of dispersant content on viscosity of pastes under various shear rates.7814) s . a P M12101 vol.% 2 vol.% 3 vol.% 4 vol.% 5 vol.% 6 vol.%(8y t i s o c s i V6420050100 Shear Rate s-1150200Figure 6: Effect of binder content on paste viscosity.2. Fabrication of solid 3D componentsCylinders and solid cones were used to demonstrate the early stages of this developing technology (Figure 7) because they are relatively simple shapes to fabricate by SFF. The defect on the cylinder surface (Figure 7) is caused by an air bubble in the paste.

The defect on the cylinder surface (Figure 7) is caused by an air bubble in the paste. The cylinder was built using a 580 µm diameter nozzle and the solid cone was built up using a 190 µm nozzle. In the current study, the FEF setup was not housed in a freezer due to limitations on the size of available freezers. Thus, the components in Figure 7 were built at room temperature. The rough surface in the photographs is caused by the collapse of deposited material as the component increases in size, and is due to instability of previously deposited materials. In future studies, components will be built in a freezer, resolving these issues.3. Fabrication of hollow conesIn order to build hollow cone geometry, a recursive offset type of algorithm was implemented. The recursive offset algorithm moves in a trajectory that is defined by the part boundaries. More in depth information about the recursive offset algorithm can be found in [14,15].

The recursive offset algorithm moves in a trajectory that is defined by the part boundaries. More in depth information about the recursive offset algorithm can be found in [14,15]. The trajectory moves in a path along the outermost boundary first. For the hollow cone example, a counter clockwise circular movement is made. After the deposition of the boundary contour, the motion path is reduced by shifting the X axis inward by a predefined amount. This shifting value is defined by the material deposition thickness. This shifting process is repeated until the innermost boundary is reached. The innermost boundary is calculated from the wall thickness value. After the innermost boundary is deposited, a straight line vector move is performed in the XZ plane. The X axis shifts inwards a predefined amount that corresponds to the desired layer overhang value while the Z elevator moves upward by a fixed amount depending on the material deposition height.

The X axis shifts inwards a predefined amount that corresponds to the desired layer overhang value while the Z elevator moves upward by a fixed amount depending on the material deposition height. The motion is then repeated in a similar fashion for the next layer. A counterclockwise circular movement is then performed, depositing the innermost boundary followed by the X axis straight line shift with this motion repeated until the new outer boundary is reached. This combination of movements is repeated until the apex of79the cone is reached. After the top layer is finished the extrusion is stopped, the Z axis is raised, and the XY table moves away from the extrusion nozzle allowing the part to be removed. Figure 8 shows the motion for the first two layers of deposition.Figure 7. FEF method fabricated cylinder and solid cone.Figure 8: Motion path for the first two layers of a hollow cone4.

Figure 8 shows the motion for the first two layers of deposition.Figure 7. FEF method fabricated cylinder and solid cone.Figure 8: Motion path for the first two layers of a hollow cone4. Freeze drying, TGA, and binder removal80Successful freeze drying of highly loaded ceramic components depends on many factors. First, it depends on the vapor pressure of water at its freezing temperature. Second, it depends on the pressure of the freeze dryer. The rate of drying also depends on the geometry and surface area of the sample. Two cone samples were produced for preliminary freeze drying tests. The samples were taken directly from the deposition table, weighed immediately, and placed in the freeze dryer for two days at -16 °C under a mechanical vacuum (∼1Pa).Table 2: Freeze drying resultsSample No.

Wet Weight (g)Dry Weight (g) Weight Loss (%)12.802.665.0022.412.266.22From the freeze drying results (Table 2), it can be seen that the mechanical vacuum pump could not meet the needs of freeze drying. Only about 30% of the water in the samples was evaporated. However, once the sample was partially freeze dried they possessed enough strength to be dried at ambient and/or elevated temperatures outside of the freeze drier. After drying, the samples were ready for binder removal. TGA results (Figure 9) show that weight loss in the samples occurs from room temperature to 600 °C in three stages. The first stage is room temperature through 120 °C. In this stage, any remaining, residual water and low melting point additives are removed. The second stage is from 120 °C to 170 °C, where glycerol is removed.

In this stage, any remaining, residual water and low melting point additives are removed. The second stage is from 120 °C to 170 °C, where glycerol is removed. The third stage is from 170 °C to 375 °C, where the higher molecular weight binders are pyrolyzed.Figure 9: TGA results from a freeze dried sample.Since the pastes in this process contain a minimum of high molecular weight organic binders, the binder removal process is relatively straightforward and can be accomplished with a81rapid heating cycle. For small samples, maximum thickness <10 mm, the binder removal can be performed using a 0.5 °C/min heating rate from room temperature to 600 °C and holding for 2 hours. The binder is completely removed under these conditions and the samples maintain their original shape.5. Sintering and sample densityThe schedule used for sintering the Al2O3 samples in the current study was 5 °C/min ramp to 1550 °C and a hold of 2 hours.

Sintering and sample densityThe schedule used for sintering the Al2O3 samples in the current study was 5 °C/min ramp to 1550 °C and a hold of 2 hours. Several sintered hollow cone samples are shown in Figure 10. Archimedes density results for the samples (Table 3) show that the samples achieved an average of 90% of their theoretical density. The fairly low density is caused by under filling, resulting in gaps between the individual filaments, during material deposition and is discussed in more detail in the next section.123Figure 10: Optical images showing the side and bottom views of sintered hollow cones fabricated by FEF.Table 3: Density test resultsSample No. Density (g/cm3)Percent of Theoretical Density13.639123.619033.5990Average3.61906. Microstructural analysis82Figure 12 and Figure 13 are SEM images obtained from area a and area b, respectively, as shown in Figure 4.

Microstructural analysis82Figure 12 and Figure 13 are SEM images obtained from area a and area b, respectively, as shown in Figure 4. The SEM images clearly indicate that the density is not uniform from area a to area b. In area a, the density is relatively high, containing only small pores having a uniform distribution. However, in area b, the density is relatively low. Several large voids are distributed throughout the area that follows a specific pattern. According to the void distribution pattern, it is surmised that under filling of the extruded filaments occurred in this area during deposition. The under filling is caused by the mismatch of XY table velocity with the extrusion rate.Figure 12: SEM images taken from area a (Figure 4) from a sectioned Al2O3 hollow cone produced using FEF.Figure 13: SEM Images taken from area b (Figure 4) from a sectioned Al2O3 hollow cone produced using FEF.IV.

Conclusions83A new SFF technique called Freeze-form Extrusion Fabrication (FEF), has been developed utilizing aqueous ceramic pastes as the build material. This is an effective, time saving, and environmentally friendly solid freeform fabrication technique. The concept was tested through successful fabrication of basic 3D geometries. Aqueous pastes consisting of a 50 ∼ 55 vol.% solids loading of Al2O3 ceramic powder, with a minimum of organic binder content, showed favorable extrusion behavior. The dispersant, binder content and pH value strongly affected the viscosity of pastes and provided control over pasted extrusion behavior. The FEF system was described in detail and major process parameters were defined. The ability of the process to build overhanging features without the use of support material was verified through fabrication of hollow cones.AcknowledgementThis work was supported by the Air Force Research Laboratory under Contract FA8650-04-C-5704References1.

EEZE-FORM EXTRUSION FABRICATION OF COMPOSITE STRUCTURESMing C. Leu1, Lie Tang1, Brad Deuser1, Robert G. Landers1, Gregory E. Hilmas2, Shi Zhang2, and Jeremy Watts21Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409 2Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractA Freeze-form Extrusion Fabrication (FEF) process capable of making three-dimensional (3D) parts and structures with graded composite materials is presented in this paper. The process development includes the design and manufacture of a gantry machine with a triple-extruder mechanism and the associated electronics hardware and computer software for fabricating functionally graded parts from multiple aqueous pastes. A rheological behavior study with Al2O3 paste is performed to identify an efficient binder for transforming the paste into a pseudoplastic with a high yield stress.

A rheological behavior study with Al2O3 paste is performed to identify an efficient binder for transforming the paste into a pseudoplastic with a high yield stress. A green part is first fabricated using the triple-extruder FEF machine in a layer-by-layer manner with the desired material gradients. The green part is then freeze-dried, its binder removed through a burnout process to obtain a brown part, and the final part obtained by sintering. The final part is analyzed using energy dispersive X-ray spectroscopy (EDS) to determine its material composition. The results demonstrate that the FEF process can be used to fabricate functionally graded composite parts with pre-specified gradients.1. IntroductionCeramic components are used increasingly in aerospace, automotive and other industrial applications due to their high heat resistance and material hardness.

IntroductionCeramic components are used increasingly in aerospace, automotive and other industrial applications due to their high heat resistance and material hardness. Traditional ceramic manufacturing processes are time consuming and expensive, especially for components with complex geometries. In recent years, several additive manufacturing processes have been developed to fabricate 3D ceramic components from CAD models in a layer-by-layer manner. They include Fused Deposition of Ceramics [1], Fused Deposition Modeling [2], Extrusion Freeform Fabrication [3], 3D Printing [4], Selective Laser Sintering [5, 6], Shape Deposition Manufacturing [7], and Robocasting [8, 9].Freeze-form Extrusion Fabrication (FEF) is a novel, environmentally friendly, additive manufacturing process that builds a 3D part layer-by-layer by computer controlled extrusion and deposition of aqueous based colloidal pastes.

Unlike most other extrusion freeform fabrication processes, which use organic binders to bond ceramic powders together, the organic binder content is only 2-4 vol% in this process, while the solids loading of the paste is up to 50 vol%. Unlike robocasting, FEF builds a “green” (before post-processing) part in an environment below the freezing point of water so as to solidify the paste after paste deposition on each layer during the fabrication process. This enables relatively large parts to be built compared with robocasting.Cones and other components from monolithic ceramics, including alumina (Al2O3) and zirconium diboride (ZrB2), have been fabricated using the FEF process; see Figure 1 [10-13].

This process has also been investigated for the fabrication of 13-93 bioactive glass scaffolds with pre-designed porosity and pore architecture [14].Figure 1: Sintered ceramic samples fabricated using the FEF process.Some key components in aerospace applications demand extremely high performance, such as the leading edges of hypersonic vehicles, missile nose cones, and nozzle throat inserts for spacecraft propulsion systems. These components must be able to withstand extremely high temperatures (> 2000°C) and integrate into the underlying substructures, which often are made of metals such as aluminum or titanium. To achieve these demanding characteristics, one approach is to build these components with functionally graded materials (FGMs), grading from an ultra high temperature ceramic (UHTC) to a refractory metal.

This grading should be done in a gradual fashion so as to minimize the thermal stresses generated due to different thermal expansion coefficients between the different materials, both during the part fabrication process and when the part is in service. Additive manufacturing processes are advantageous for building such components with functionally graded materials.This paper describes our design and development of a triple-extruder FEF system for fabricating parts with functionally graded materials. The part fabrication process by this system involves computer control of flows of multiple aqueous pastes (each controlled separately), the mixing of these pastes, and the extrusion of the mixed paste to fabricate a 3D part layer-by-layer according to a CAD model and pre-specified material compositions. The system development also includes planning and control algorithms and software for motion control and extrusion control.

The system development also includes planning and control algorithms and software for motion control and extrusion control. This paper details the process and system concept, system design and development, and some experimental results with the developed system.2. Process and system conceptThe specific aim of our development of the FEF system is to extrude three different aqueous pastes (made of WC+ZrO2, W, and ZrC) at controlled rates to build parts having a gradient from ZrC to W, as pre-specified, after binder burnout and reactive sintering of the green part built by the FEF system. Figure 2 shows the concept of fabricating a part with gradient materials using the system. As shown in the figure, a green part is first fabricated by the system in a layer-by-layer manner. The fabricated green part is then freeze-dried, and the binder is removed through a burnout process to produce a “brown” (binder pyrolyzed) part.

The fabricated green part is then freeze-dried, and the binder is removed through a burnout process to produce a “brown” (binder pyrolyzed) part. The brownpart then undergoes reactive sintering to obtain the final part, which has CAD dimensions and FGM compositions that are pre-specified.Figure 2: Procedure for fabricating gradient materials using the FEF process.Figure 3: Triple-extruder mechanism design.To realize the process of fabricating an FGM part, we have designed and developed an FEF machine equipped with three servo controlled extruders and a paste mixer. Fabricating an FGM part that grades between ZrC and W after sintering requires three pastes: ZrO2+WC, W, and ZrC. The reactive sintering process involves the following chemical reaction:ZrO2 + 3WC = ZrC + 3W + 2COThe three pastes are driven simultaneously by a triple-extruder mechanism, as illustrated in Figure 3.

Continuous control over the material compositions and their gradients during the part building process can be achieved by controlling and coordinating the flow rates for the different pastes. As an example, assuming that the three barrels containing the three different pastes have the same cross-sectional area, a desired paste mixture consisting of 20% paste A, 30% paste B, and 50% paste C can be achieved by controlling the three plunger velocities with the ratios of V1:V2:V3 = 2:3:5, where V1, V2, and V3 are the plunger velocities for pastes A, B and C, respectively. Note that the volumetric paste flow rate is equal to the plunger velocity multiplied by the cross-sectional area of the plunger. The different paste flow rates, together with the percentage of solids loading of each paste and also the material density and molar mass data available from the literature, allows for calculation of the material compositions of the final part.

For example, assume that paste A is ZrO2+WC, paste B is W, and paste C is ZrC, and that the percentage solids loading is the same for the three different pastes. From the above ratios of pastes A, B and C, the weight of ZrC from paste B and the weight of W from paste C can be determined. The ratio of ZrO2 and WC in paste A should be such that both ZrO2 and WC are fully consumed when they are converted to ZrC and W during the reactive sintering process. This allows for calculation of the weights of both ZrO2 and WC for a given amount of paste A, and from these weights, the weights of ZrC and W after reactive sintering can be determined. By adding these ZrC and W weights to the weight of ZrC from paste B and the weight of W from paste C, the ratio of ZrC and W in the final part can be obtained. Using the procedure described above, the weight percentage of ZrC vs. W is determined to be 17.3% vs. 82.7% for the given example.3.

Using the procedure described above, the weight percentage of ZrC vs. W is determined to be 17.3% vs. 82.7% for the given example.3. Triple-extruder FEF system design and developmentThe triple-extruder mechanism was designed using three stainless steel barrels, each containing a paste driven by a plunger whose movement is controlled by a DC servo motor (Kollmorgen AKM23D); see Figure 4. The encoder signal from the servo amplifier provides a resolution of 0.62 µm for the plunger’s movement. The flow rate of paste in each barrel is controlled by the plunger’s velocity, and the force exerted on the plunger is measured by a load cell (Omega LC-305). A key component of the triple-extruder mechanism is a mixer connected to the ends of the three barrels.

A key component of the triple-extruder mechanism is a mixer connected to the ends of the three barrels. The FEF system uses a static mixer, which merges the three different pastes and mixes them into a homogeneous stream as the combined pastes pass a series of mixing blades positioned at alternating angles.The triple-extruder mechanism is mounted on a gantry system, which consists of three orthogonal linear drives (Velmex, Bloomfield, NY), each with a 508 mm travel range. The X- axis consists of two parallel slides and is used to support the Y-axis. The use of two parallel slides (vs. a single slide) provides a smoother and more stable motion, thus providing a larger work space for part fabrication. The Z-axis is mounted on the Y-axis, and the extrusion mechanism is mounted on the Z-axis. Each of these axes is mounted with limit switches on both ends.

The Z-axis is mounted on the Y-axis, and the extrusion mechanism is mounted on the Z-axis. Each of these axes is mounted with limit switches on both ends. Four DC servo motors (Pacific Scientific PMA22B), each with a resolver for position feedback at a resolution of 1000 counts per revolution, drive the various axes. Each motion axis has a maximum speed of 127 mm/s and a position sensor resolution of 0.00254 mm.Figure 4: FEF machine with triple-extruder mechanism.The part fabrication process is carried out in a freezing environment, which can be controlled to as low as -20°C using a liquid nitrogen injection system. This enables the aqueous paste to solidify at temperatures below the freezing point of water after it is extruded. The paste solidification helps avoid part deformation during the fabrication process, thus enabling fabrication of larger parts.

The paste solidification helps avoid part deformation during the fabrication process, thus enabling fabrication of larger parts. To keep the paste from freezing before its exit from the deposition nozzle, a heating jacket is used to keep the paste’s temperature above the water freezing point while it is inside the barrels and the mixer. In our present study, the freezer’s temperature is kept at -10°C, while the heating jacket’s temperature is kept at 10°C.The FEF control system is developed with a personal computer that has a Turbo PMAC board from Delta Tau Motion System, as shown in Figure 5. The 3-axis movement of the gantry system is controlled by a motion control program with the Delta Tau Turbo PMAC PCI board. The control software is written using the Delta Tau language and downloaded to the Turbo PMAC board via PEWIN32, which is the Delta Tau programming interface.

The control software is written using the Delta Tau language and downloaded to the Turbo PMAC board via PEWIN32, which is the Delta Tau programming interface. Extrusion of pastes is controlled with three servo motors using a National Instruments PXI chassis and LabVIEW.Linear Cylinder AmplifiersA/DDelta Tau Turbo PMAC3 Axis Gantry SystemLoad CellsD/ALinear CylindersExtrusion SystemPC with PEWIN32Signal AmplifiersC/TFigure 5: The FEF control system schematic.4. Paste preparation and testingThe pastes to be extruded each must have an engineered composition and proper rheological behavior for extrusion through a fine orifice to produce a 3D part.

Paste preparation and testingThe pastes to be extruded each must have an engineered composition and proper rheological behavior for extrusion through a fine orifice to produce a 3D part. The objective of paste preparation is to develop paste recipes that provide high solids loading (>45 vol% to as high as 60 vol% if possible), no phase separation (i.e., liquid phase migration) under pressure, homogeneity of the paste throughout the extrusion process, and pseudoplastic behavior with a defined yield stress. To limit the difference in sintering shrinkage along the cross section of FGM components, WC+ZrO2, W and ZrC were selected as the three pastes. The raw powders (W, ZrO2, and WC) are characterized using surface areas and particle sizes.In order to obtain homogeneous aqueous pastes, the raw powders must be well dispersed in water. To this end, a series of settling tests were performed using Darvan C as the surfactant at various pH values in water.

To this end, a series of settling tests were performed using Darvan C as the surfactant at various pH values in water. The optimal conditions for the ZrO2, W and WC powders were determined to be 0 mg per gram of ZrO2 in H2O at pH=9, 2 mg per gram of W in H2O at pH=7, and 0 mg per gram of WC in H2O at pH=8.To study the rheological behavior of pastes under pressure, a capillary rheometer was designed and fabricated, as shown in Figure 6.

The rheological behavior and viscosity of Al2O3 slurries containing 40 vol%, 45 vol% and 50 vol% solids loading were determined using the capillary rheometer under various forces based on the following formulae for shear stress, shear rate, and viscosity:LPRsShearStres2:34:RQShearRate4:8PRViscosityLQFigure 6 The capillary rheometer.where P is pressure, R is nozzle radius, L is nozzle length, and Q is volumetric paste flow rate.The measured results (Figure 7) show that all of the pastes exhibited pseudoplastic behavior with no visible phase separation. Thus, the Herschel-Buckley model can be used to model and analyze the pastes. The 50 vol% Al2O3 paste has the highest yield stress (~60 Pa) among the three tested pastes.Figure 7: Rheological behavior of Al2O3 paste with different solid loadings.An appropriate organic binder is required to provide green body strength and to achieve a desirable rheological behavior for extrusion.

Three types of water-soluble binders (PVA, Aquazol 200, and Methycell) were examined using Al2O3 as a replacement powder for the ZrO2, W and WC powders in order to minimize experimentation costs. The viscosity and rheological behavior of a series of Al2O3 slurries with various quantities of organic binders were studied. The results indicated that Methycell is the most efficient binder among the three for transforming the rheological behavior of Al2O3 slurries into a pseudoplastic with a high yield stress. As shown in Figure 8, a 0.5 wt% Methylcell addition allows an Al2O3 slurry containing 45 vol% solids to exhibit a yield stress of 220 Pa. In contrast, as shown in Figure 9, an Al2O3 slurry containing 45 vol% solids with a 1.0 wt% Aquazol 200 addition exhibits a yield stress of 130 Pa.Figure 8: Effect of Methylcell on the Al2O3 paste’s (45% solid loading) rheological behavior.Figure 9: Effect of Aquazol 200 on the Al2O3 paste’s (45% solid loading) rheological behavior.5.

Results of system evaluationTo test the capabilities of the FEF system for building parts with gradient materials, a cylindrical part with a 50 mm diameter was fabricated with two extruders filled with limestone (CaCO3) pastes, one in green color and the other in pink color. The fabrication result is shown in Figure 10. The color of the fabricated cylinder part starts in pink (A) and shifts to brown (B), then green (C), then brown (D), then pink (E), and finally green (F). The color distribution of the part is consistent with the velocity profiles of the two extruders shown in Figure 11. In section A, only extruder 1 was moving, which resulted in the pink color. When both extruders are controlled to move at the same velocity, the extrudate becomes light brown, as shown in section B. Figure 11 also shows that the extrusion force for both extruders remains in the range of 300 to 400 N.

Figure 11 also shows that the extrusion force for both extruders remains in the range of 300 to 400 N. Figures 10 and 11 demonstrate that a part can be built with material gradients by varying extrusion velocities for different pastes. Note that the planning and control algorithms must take into consideration a delay between the time the extrusion material is switched to the time the switched material exits the nozzle.After the visual validation, an investigation of the capability of the FEF system to build an FGM part by varying the ratios of two pastes consisting of different materials, one pure alumina (Al2O3) and the other 50Vol% Al2O3 + 50Vol% ZrO2, was performed. For examining the material compositions of the built part, pre-mixed compositions of 50% Al2O3 + 50% ZrO2, 75% Al2O3 + 25% ZrO2, and 100% Al2O3 were lab-prepared into pellets, freeze-dried, and sintered as a control set.

They then were coated with gold palladium in preparation for energy dispersive X-ray spectroscopy (EDS). The EDS results are shown in Figure 12 for 50% Al2O3 + 50% ZrO2 and in Figure 13 for 75% Al2O3 + 25% ZrO2. Figure 14 shows the calculated ratio of the peak height of zirconium (Zr) vs. the peak height of aluminum (Al) for the pellets made of the three different pastes. These ratios were then used as the basis to examine the material compositions of parts fabricated by the FEF system.Figure 10: Fabrication of cylindrical part with discrete material gradients.Figure 11: Control signal and measured extrusion force and velocity profile.Figure 12: EDS intensity distribution of lab-prepared specimens with 50% Al2O3 + 50% ZrO2.Figure 13: EDS intensity distribution of lab-prepared specimens with 75% Al2O3 + 25% ZrO2.01020304050607002550EDS Intensity Ratio of Zr to Al (%)Composition Ratio of ZrO2to Al2O3(%)Figure 14: Ratio of Zr peak height vs.

Al peak height from EDS results for three test pellets.Parts fabricated using the FEF system began with 100% Al2O3 paste, transitioned to one half of 100% Al2O3 paste and one half of 50% Al2O3+50% ZrO2 paste (thus, 75% Al2O3+25% ZrO2 in overall composition) and ended with 50% Al2O3+50% ZrO2 paste in building a test bar. Once fabricated, the Al2O3-ZrO2 parts were freeze-dried and sintered to a final height of 32 mm. The freeze-drying was carried out at a temperature of -25 °C and a pressure of about 3000 Pa. The sample was held at this temperature and pressure for 24 hours. The sample was then allowed to warm to room temperature (25 °C) while maintaining the same vacuum pressure and held for an additional 24 hours. Binder burnout was accomplished by heating the samples at 1 °C/min up to 600 °C and holding for 1 hour. From there, the samples were heated at 10 °C/min up to 1550 °C and held for 90 minutes.

Binder burnout was accomplished by heating the samples at 1 °C/min up to 600 °C and holding for 1 hour. From there, the samples were heated at 10 °C/min up to 1550 °C and held for 90 minutes. Following the 90-minute hold, the samples were cooled back to room temperature at 25 °C/min.The sintered part was then cut, polished, and coated with gold palladium in preparation for EDS measurements. The results are shown in Figure 15 for average EDS intensity measurements over 8 sections with an area 4 mm wide by 3 mm high per section. Figure 16 plots the ratio of Zr peak height vs. Al peak height calculated for each of these sections. For the control test pellets, the ratio of Zr peak height vs. Al peak height was calculated to be 62% (Figure 14) for the composition of 50% Al2O3+50% ZrO2, compared to the FEF fabricated part, which achieved a ratio of 55% for the same composition (Figure 16).

Al peak height was calculated to be 62% (Figure 14) for the composition of 50% Al2O3+50% ZrO2, compared to the FEF fabricated part, which achieved a ratio of 55% for the same composition (Figure 16). This slight discrepancy could be due to the width of the scan area edging into the transition region to the next composition in the FEF built part. Because the intensity measurement is taken as an average over the area, and the edge of the area could contain part of the 75% Al2O3 + 25% ZrO2 mixture, the respective measurement of Zr/Al intensity ratio would decrease slightly. As another comparison, the 75% Al2O3 + 25% ZrO2 test pellets measured a ratio of 20% Zr/Al, and the FEF fabricated part was calculated to be 20% as well.

As another comparison, the 75% Al2O3 + 25% ZrO2 test pellets measured a ratio of 20% Zr/Al, and the FEF fabricated part was calculated to be 20% as well. This compares very well because the scan area was contained fully within the desired region.(a)(b)(c)(d)(e)(f)(g)(h)Figure 15: EDS intensity distribution measured for different sections of an FEF fabricated part with varying ratios of Al2O3 and ZrO2 along the measurement direction.75% Al2O325% ZrO250% Al2O350% ZrO201020304050600510152025303540Ratio of Zr to Al (%)Position (mm)100% Al2O3Figure 16: EDS intensity measurements for FEF fabricated test bar.6. ConclusionsA Freeze-form Extrusion Fabrication (FEF) process aimed at fabricating 3D parts that grade between ultra high temperature ceramics and refractory metals after reactive sintering is presented in this paper.

The main process concept is to mix multiple aqueous pastes according to part material composition requirements and to extrude the mixed paste to fabricate a 3D part layer-by-layer in an environment below water’s freezing temperature. Based on this concept, a triple-extruder FEF system including the mechanical machine, electronics hardware, and computer software has been developed. For paste preparation, a rheological behavior study using Al2O3 paste indicates that Methycell is an efficient binder for transforming the paste’s rheological behavior into a pseudoplastic with a high yield stress. The capability of the developed FEF system for fabricating 3D parts with desired material gradients is validated first by observing the color transitions and the corresponding extrusion velocity profiles in a fabricated limestone part and then by measuring energy dispersive spectroscopy (EDS) peak intensities in a fabricated test bar with varying Al2O3/ZrO2 ratios along the measurement direction.

In the future, investigations using the developed FEF system to fabricate parts with the three pastes: ZrO2+WC, W, and ZrC, followed by post-processing to produce parts that grade between ZrC and W after reactive sintering, will be performed. The sintered specimens will be evaluated by measuring density, material compositions, and mechanical properties and by characterizing microstructures.7. AcknowledgementsThis project is funded by NSF grant #CMMI-0856419 with matching support from Boeing Company through the Center for Aerospace Manufacturing Technologies at the MissouriUniversity of Science and Technology, and by the Air Force Research Laboratory through Universal Technology Corporation (Contract #10-S568-0094-01-C1). The technical support of Samuel Easley and Michael Hayes of Boeing Research & Technology is greatly appreciated.8. References[1] A. Bandyopadhyay, P. Panda, M. Agarwala, S. Danforth, and A.

eeze-form Extrusion Fabrication of Functionally Graded Material Composites UsingZirconium Carbide and TungstenAng Li, Aaron S. Thornton, Bradley Deuser, Jeremy L. Watts, Ming C. Leu, Gregory E. Hilmas, Robert G. Landers Missouri University of Science and Technology, Rolla, MO 65409AbstractUltra-high-temperature ceramics are being investigated for future use in aerospace applications due to their superior thermo-mechanical properties, as well as their oxidation resistance, at temperatures above 2000°C. However, their brittleness makes them susceptible to thermal shock failure. As graded composites, components fabricated as functionally-graded materials (FGMs) can combine the superior properties of ceramics with the toughness of an underlying refractory metal. This paper discusses the grading of two materials through the use of a Freeze-form Extrusion Fabrication (FEF) system to build FGM parts consisting of zirconium carbide (ZrC) and tungsten (W).

This paper discusses the grading of two materials through the use of a Freeze-form Extrusion Fabrication (FEF) system to build FGM parts consisting of zirconium carbide (ZrC) and tungsten (W). Aqueous-based colloidal suspensions of ZrC and W were developed and utilized in the FEF process to fabricate test bars graded from 100%ZrC to 50%W- 50%ZrC (volume percent). After FEF processing, the test bars were co-sintered at 2300°C and characterized to determine their resulting density and microstructure. Four-point bending tests were performed to assess the flexural strength of the test bars made using the FEF process, compared to that prepared using conventional powder processing and isostatic pressing techniques, for five distinct ZrC-W compositions. Scanning electron microscopy (SEM) was used to examine the inner structure of composite parts built using the FEF process.1.

Scanning electron microscopy (SEM) was used to examine the inner structure of composite parts built using the FEF process.1. IntroductionFabrication of a functionally graded material (FGM) part refers to the process of manufacturing a part with multiple materials in a graded fashion in order to take advantage of complementary material properties while minimizing residual stresses that may result from the sintering process [1]. Ceramics are often used in high-temperature applications for their superior heat resistance; however, poor fracture toughness limits their use in high-stress scenarios and they are often difficult to manufacture for complex geometries using traditional processes.

Several additive manufacturing technologies have been developed in recent years that can fabricate ceramic components with complex geometries, but few have the ability to build FGM parts.Additive manufacturing (AM) technology has evolved since its inception in the mid 1980’s, from polymer-based processes to metal-based and ceramic-based processes. Stereolithography [2], Fused Deposition Modeling [3], 3D Printing [4], and Selective Laser Sintering [5,6] are among the popular AM technologies practiced in industry today. The current metal and ceramic AM technologies are mostly limited to single material (monolithic) part fabrication.

The current metal and ceramic AM technologies are mostly limited to single material (monolithic) part fabrication. Robocasting [7], Extrusion Freeform Fabrication [8], Shape Deposition Manufacturing [9], and Laser Metal Deposition [10] are more apt in their potential to building multiple-material parts since they are deposition-based processes.467This study considers a novel additive manufacturing technology called Freeze-form Extrusion Fabrication (FEF), which is capable of producing three-dimensional parts by depositing aqueous-based ceramic and metal pastes in a layer-by-layer manner within a sub-zero temperature environment to minimize the amount of organic binder necessary, thus making post- processing easier and more environmentally friendly [11-15].Some key components in aerospace applications demand extremely high performance, such as the leading edges of hypersonic vehicles, missile nose cones, and nozzle throat inserts for spacecraft propulsion systems.

These components must be able to withstand extremely high temperatures (> 2000°C) and be integrated with underlying substructures, which are typically made of metals such as aluminum or titanium. To achieve these demanding characteristics, one approach is to build these components while grading from a ceramic to a metal. The grading should be done in a gradual fashion so as to minimize the thermal stresses generated due to different thermal expansion coefficients between the different materials, both during part fabrication and when the part is in service. Deposition-based additive manufacturing processes are advantageous for fabricating such components with functionally graded materials.This paper considers the sintering of ZrC-W composite parts fabricated by the FEF process.

As a refractory metal with a melting point of 3422°C and yield strength of approximately 800 MPa at room temperature[17], W shows great potential in aerospace applications, such as re- entry components, uncooled liquid rocket chambers, rocket nozzles, electrical propulsion components, etc. However, dramatic decreases occur in the mechanical strength of W with increases in temperature. For example, the mechanical strength of W decreases by ~60% when heated from room temperature to 1000°C [17]. To enhance the mechanical strength of W at elevated temperatures, ZrC can be introduced as a reinforcement material because of its high melting temperature (3532°C) and thermal expansion coefficient similar to that of W [18]. Several methods have been used for the fabrication of ZrC-W composites, the most typical one being hot pressing [18].

Several methods have been used for the fabrication of ZrC-W composites, the most typical one being hot pressing [18]. A high relative density can be achieved by the hot pressing method (＞ 98%); however, the dimensional limitation narrows the scope of applications for this method. An in-situ reaction sintering process was able to achieve a relative density of approximately 94.5% for the manufacture of ZrC-W composites without applying external pressure during sintering [19]. In this paper, we compare the relative densities, flexural strengths, and microstructures of test bars of different ZrC and W compositions fabricated by the FEF process and by a traditional powder processing route that includes isostatic pressing. The test bars were sintered at temperatures ranging from 2100 to 2300°C in attempts to achieve the highest possible relative densities.2.

The test bars were sintered at temperatures ranging from 2100 to 2300°C in attempts to achieve the highest possible relative densities.2. Freeze Form Extrusion Fabrication Process OverviewAn FEF machine equipped with three servo controlled extruders and a three-axis gantry motion system in a temperature-controlled enclosure has been developed, as shown in Figure 1. The FEF machine is equipped with a triple-extruder mechanism to build three-dimensional FGM parts with complex geometries. The different materials are combined by feeding them into an inline static mixing unit and output the mixed pastes through a single orifice. This static mixer forces the pastes to mix together before exiting the orifice, and provides a natural transition between paste composition changes. The different pastes are extruded simultaneously by controlling the velocity of each plunger.

The different pastes are extruded simultaneously by controlling the velocity of each plunger. As an example, assuming that the three cylinders contain three different pastes and have the same cross-sectional area, a desired paste mixture468consisting of 20% paste A, 30% paste B, and 50% paste C can be achieved by controlling the three plunger velocities with the ratios of v1:v2:v3 = 2:3:5, where v1, v2, and v3 are the plunger velocities for pastes A, B, and C, respectively. The mixed colloidal paste is deposited onto a solid substrate layer by layer within a sub-zero temperature environment (-10°C in our present study). Following the FEF fabrication, the fabricatedpart is transferred to a freeze-dryer to sublime excess water from the green part and then undergoes binder burnout and sintering.Figure 1. The triple-extruder Freeze-form Extrusion Fabrication (FEF) machine.3. Materials Processing3.1.

The triple-extruder Freeze-form Extrusion Fabrication (FEF) machine.3. Materials Processing3.1. Characterization of Sintering BehaviorIn order to achieve acceptable mechanical properties, ZrC-W composites produced by the FEF process will require a high temperature sintering cycle to achieve the highest possible relative density. The co-sinterability of ZrC and W was investigated by fabricating several batches of test pellets with the composition of 50vol%ZrC+50vol% W and sintering at different temperatures and heating rates (as shown in Table 1).Table 1. Heating conditions for co-sintering test.Temperature Holding time Atmosphere 1 hour 1 hour 3 hours 3 hours2100°C 2300°C 2300°C 2300°CHelium Helium Helium HeliumHeating rate 10 °C/min 10 °C/min 10 °C/min 10°C/min from room temperature to 2100°C, and then 2°C/min to 2300°CRelative density 74.51% 71.52% 80.71% 78.45%During the experiments, the powders of ZrC (＜2μm, Grade B, H.C.

Starck, Karlsruhe, Germany) and W (0.6~1μm, Sigma Aldrich, St. Louis, MO) were first mixed and ball-milled using acetone and zirconia media for 2 hours. After ball milling, the slurry was dried by rotary evaporation (Buchi, Flawil, Germany) at a temperature of 70°C, low vacuum (~27 kPa), and a469rotation speed of 60 rpm. The dry powder was ground, filtered through a 180μm sieve, and pressed into pellets using a hydraulic press with a half-inch diameter die at 2000 psi. The pellets were then isostatically pressed at 30,000 psi before sintering. Sintering was performed in a graphite furnace (Thermal Technology, Santa Rosa, CA) under a helium atmosphere. The densities of the sintered pellets were determined using the Archimedes method, and each pellet was polished for scanning electron microscopy (SEM, S4700 and S570, Hitachi, Tokyo, Japan).

The densities of the sintered pellets were determined using the Archimedes method, and each pellet was polished for scanning electron microscopy (SEM, S4700 and S570, Hitachi, Tokyo, Japan). Figure 2 shows the pressed and sintered pellets.Figure 2, Pressed and sintered pellet.The sintering of the pellets (co-sintering of ZrC and W) was performed under four different heating conditions, which are listed along with the resulting relative density data in Table 1. Maintaining a heating rate of 10°C/min while increasing the sintering temperature from 2100°C to 2300°C was found to increase the relative density from ~71% to ~81% (see Table 1). However, a relative density above 81% could not be achieved despite additional modifications to the sintering cycle. SEM analysis of a sample with the highest relatively density (~81%) showed many visible pores in the ZrC phase as seen in Figure 3, where ZrC is the darker phase and W is the lighter phase.

SEM analysis of a sample with the highest relatively density (~81%) showed many visible pores in the ZrC phase as seen in Figure 3, where ZrC is the darker phase and W is the lighter phase. This indicates that the rate of grain growth of ZrC was too fast under the current heating condition, causing entrapped pores and thus insufficient densification of ZrC.Figure 3. SEM images of 50%ZrC+50%W pellets with relative density of 80.71%.470Table 2. Comparison between as-received and attrition-milled ZrC powders.ZrC powder As-received Attrition-milledAverage particle size (μm) 0.47 0.35BET surface area (m2/g) 1.06 6.13To verify the effect of attrition milling on the sinterability of ZrC, another batch of pellets was made with the composition of 50vol%ZrC+50vol%W and 90vol%ZrC+10vol%W using the attrition milled ZrC powder. Also, one batch of 90vol%ZrC(as-received)+10vol%W pellets were also produced for comparison.

Also, one batch of 90vol%ZrC(as-received)+10vol%W pellets were also produced for comparison. The sintering test was conducted at a heating rate of 10°C/min from room temperature to 2100°C, and then 2°C/min to 2300°C in order to allow for more time at the lower sintering temperatures and limit ZrC grain growth.The pellets made with the attrition-milled ZrC powder resulted in a significant increase in density compared with the pellets made with the as-received ZrC powder (Table 3). The 50vol%ZrC+50vol%W composition achieved a 98.9% the 90vol%ZrC+10vol%W composition achieved a relative density of 99.4%. Figure 4 shows SEM images for 50vol%ZrC+50vol%W pellets made with the as-received and attrition-milled ZrC powders. Based on SEM analysis, it is clear that attrition milling not only contributed to a higher material density after sintering, but also decreased the ZrC and W grain sizes as well as the number and size of pores inside the sintered parts.

Thus, the sintered pellets made with attrition- milled ZrC powder resulted in a higher density and finer microstructure.relative density, whileTable 3. Comparison between sintered pellets made with as-received and attrition-milled ZrC powders.Composition50vol%ZrC+50vol%W 90vol%ZrC+10vol%WRelative density (with as- received ZrC) 78.45% 71.32%Relative density (with attrition- milled ZrC) 98.9% 99.4%471(a) (b)Figure 4. (a). Microstructure of 50vol%ZrC+50vol%W pellets made from as-received ZrC powder. (b). Microstructure of 50vol%ZrC+50vol%W pellets made from attrition-milled ZrC powder.Figure 5. XRD analysis on ZrC/W pellets.The sintered ZrC/W pellets, made using the attrition-milled ZrC powder, were ground back to powder for X-ray diffraction (XRD) analysis (XDS 2000, Scintag Inc., Cupertino, CA). The XRD results are presented in Figure 5.

The XRD results are presented in Figure 5. The detected highest-intensity ZrC peaks (55.841o, 66.612o, 69.985o and 82.921o) in the XRD patterns were all shifted to higher two-theta angles compared to the reference peaks (55.633o, 66.348o, 69.712o and 82.589o). An increase in the two- theta angle indicates a decrease in lattice parameter in the ZrC unit cell, implying that there was formation of (Zr,W)C solid solution. This is not an unexpected result and has been previously observed in the technical literature for ZrC-W systems [19].3.2. Development of ZrC and W PastesTo ensure a stable viscosity during FEF deposition and good mechanical strength of parts made by the FEF process, many tests were conducted on the development of ZrC and W pastes. Most important to these tests were the investigation of appropriate dispersants, which control the rheological behavior of the pastes.

Most important to these tests were the investigation of appropriate dispersants, which control the rheological behavior of the pastes. W powder was found to be particularly difficult to disperse in aqueous solutions due to its high density (19.25 g/cm3). After an initial set of experiments472involving several dispersants, it was determined that Sodium Dodecyl Sulfate (SDS, Sigma Chemical Co, St. Louis, MO) could disperse the W powder in distilled water at a solids loading of ~45 vol%. Further, SDS also performed well in dispersing ZrC powder to ~50 vol% solids loading in distilled water. Thus SDS was selected as the dispersant to produce pure ZrC and 50vol%ZrC+50vol%W pastes at solids loadings of 50 vol%. After each powder composition was dispersed in water, hydroxypropyl methylcellulose (Methocel, The Dow Chemical Company, Midland, Michigan) was added to the slurries as the binder to control the paste viscosity.

The amount of dispersant and binder were controlled to ensure paste viscosities that suited the FEF machine. During the FEF fabrication, the viscosity of pastes was adjusted based on the extrusion force of the plunger by keeping the extrusion force in the range between 600N and 900N. The pastes were then extruded by FEF machine to fabricate test bars for flexural strength testing.4. FEF Test Bar FabricationTest bars of five different compositions were fabricated using the FEF process, among which three and 37.5vol%W+62.5vol%ZrC, were deposited by mixing of the two initial pastes (100vol%ZrC and 50vol%ZrC+50vol.%W) in various ratios. Five additional test bars were made using an isostatic press after mixing W and attrition-milled ZrC powders into the desired compositions in order to compare their mechanical properties. The process of fabricating test bars was the same as that for pellets.

The process of fabricating test bars was the same as that for pellets. After sintering, these test bars were cut and ground into 3×4×45 mm3 pieces according to ASTM C 1161-02c for type B bars.compositions,12.5vol%W+87.5vol%ZrC,75vol%W+25vol%ZrCTable 4. Measured properties of FEF fabricated and iso-pressed test bars.FEF FabricatedIsostatic PressedComposition (vol.%) 100%ZrC 12.5%W+87.5%ZrC 25%W+75%ZrC 37.5%W+62.5%ZrC 50%W+50%ZrCRelative Density 62.05% 47.89% 56.19% 47.28% 70.08%Flexural Strength (MPa) 73 25 25 28 31Relative Sintered Density 98.49% 94.41% 97.34% 95.40% 99.81%Flexural Strength(MPa) 224 265 398 414 404The data in Table 4 indicate that the flexural strength of the isostatic pressed bars increased with a higher concentration of W, from 224 MPa for 100% ZrC to 404 MPa for the 50%ZrC+50%W composition. This result implies that the W played an important role in increasing the flexural strength of the composites.

This result implies that the W played an important role in increasing the flexural strength of the composites. However, this trend was not observed for the FEF fabricated bars. The relative density and flexural strength of FEF fabricated bars were, in general, much lower than those of the isostatic pressed bars. In order to have a better understanding of the differences, SEM images were taken to compare the microstructures of the test bars. From the SEM analysis (Figure 6), many large pores (10’s to 100’s of µm) were present in the bars produced by the FEF process, and the pores became even more severe with increasing ZrC content. However, all of the isostatic pressed bars show highly densified microstructures, with the relative densities of all test bars above 94% and the 50%ZrC+50%W composition achieving 99.8% in relative density.473There are two reasons that can be used to explain for the severe porosity in the FEF fabricated bars.

The first reason is that the static mixer did not do an adequate job of mixing the pastes, leading to significant variations in the materials compositions and differential sintering at different locations within the test bars. Figure 7 shows the cross sections of FEF fabricated test bars with 62.5%ZrC+37.5%W and 87.5%ZrC+12.5%W. Clearly visible in the images are marked differences in the intended compositions at different locations. The ZrC-rich areas (darker phase) are distinct from W-rich areas (lighter phase), indicating the insufficient paste mixing during the extrusion. The second and more critical reason for the porosity is that during the FEF process large ice crystals were formed as the water was freezing. This has been shown in previous studies of aqueous based freeze casting of ceramics wherein large ice crystals are formed, resulting in large voids (100’s of µm) after sintering [20].

This has been shown in previous studies of aqueous based freeze casting of ceramics wherein large ice crystals are formed, resulting in large voids (100’s of µm) after sintering [20]. Once formed during the freezing process, these large defects remained inside the test bars, affecting the final mechanical properties. An example is given in Figure 8, which shows the formation of what appears to be the remnants of ice crystals in the 100%ZrC sample after freeze drying. As has been discussed by Sofie and Dogan [20], these voids may be controlled, or even eliminated, in the future through the use of glycerol additions to the aqueous based slurries. .(a) (b)474(c)(e) (f)(g) (h)475(d)(i)(j)Figure 6. SEM images of the micro-structure of (a). 50%ZrC+50%W, FEF fabricated; (b). 50%ZrC+50%W, isostatic pressed; c). 62.5%ZrC+37.5%W, FEF fabricated; (d). 62.5%ZrC+37.5%W, isostatic pressed; (e). 75%ZrC+25%W, FEF fabricated; (f). 75%ZrC+25%W, isostatic pressed; (g).

Summary and ConclusionsZrC/W test bars with compositions ranging from 100% ZrC to 50%ZrC+50%W by volume were prepared by isostatic pressing and freeze-form extrusion fabrication (FEF). Using attrition milled ZrC powder, relative densities ranging from 94.4% to 99.8% could be achieved for isopressed and co-sintered ZrC and W compositions by sintering at 2300°C at the heating rate of 10°C/min from room temperature to 2100°C, and then 2°C/min to 2300°C. However, FEF fabricated bars only achieved 47-70% relative density for similar compositions. Flexural strength data from four-point bending tests followed a similar trend, with strengths in the range of 200-400 MPa for the dense isopressed bars and strengths of only 25-70 MPa for the bars produced by FEF processing.

The large differences in both density and flexure strength between the isostatic pressed and FEF fabricated test bars were found to be due to large voids inside the FEF fabricated bars from the formation of ice crystals during the freezing process. Elimination of these large voids, combined with an improved mixing method, are needed in order to enhance the performance of graded ZrC/W composites fabricated by the FEF process.6. AcknowledgementsThis project is funded by NSF grant #CMMI-0856419 with matching support from the Boeing company through the Center for Aerospace Manufacturing Technologies at the Missouri University of Science and Technology, and by the Air Force Research Laboratory contract #10- S568-0094-01-C1 through the Universal Technology Corporation.4777. References[1] Kruth, J.P., Leu, M.C., Nakagawa, T., 1998, Progress in Additive Manufacturing and Rapid Prototyping, CIRP Annals - Manufacturing Technology, 47/2:525-540.

eeze-Spray Processing of Layered Ceramic CompositesQ. Fu, O. Jongprateep, A. Abbott and F. Dogan Department of Materials Science and Engineering, University of Missouri-Rolla Rolla, MO, 65409AbstractThermal gradients and associated stresses are critical in designing with ceramiccomposites having low thermal conductivity. In order to reduce the stresses from thermal gradients, compositional gradients are employed in designing of composite structures. This study addresses development of freeze-spray process to fabricate layered ceramic structures with controlled layer thickness and microstructural development. The composites were processed by spraying of ceramic slurries with low binder content and relatively high solids loadings (up to 40 vol%) on a cooled substrate. The frozen parts were freeze-dried and sintered at elevated temperatures.

The frozen parts were freeze-dried and sintered at elevated temperatures. The relationship between microstructural development and thermal expansion behavior of Al2O3 and Y2O3-stabilized ZrO2 functionally graded ceramic composites is discussed.IntroductionFunctionally graded materials are a new generation of engineered materials that havebecome of much interest in recent years. The graded materials are ideal candidates for various applications ranging from functional and structural materials. Such applications involve severe thermal gradients, such as components in advanced aerospace engines, and thermal barrier coatings for tools, to biomedical implants1-2. A gradual transition of material compositions in a solid creates gradual variations in microstructural development which may reduce the thermal expansion and sintering rate mismatch3-7.The present study describes the development of a freeze-spray process for rapidfabrication of graded lamina composites.

Freeze-spray process involves the preparation of ceramic slurries with a low binder content that are sprayed onto a cold substrate to obtain a compositional gradient. Frozen parts are then subjected to freeze drying process, so that green bodies with low binder content can be obtained 8-16. The process is also suitable when spray- freezing of non-aqueous slurries are required. Since the green bodies contain relatively low amounts of binder, a rapid binder burn-out process can be employed for large volume parts prior to the sintering. Freeze-spray process is a versatile and relatively simple technique to fabricate graded structures of various materials which may not be obtained by other methods such as lamination of green tapes processed by tape casting or thermal (plasma) spray processes. By changing slurry compositions during freeze-spray process, one can achieve gradual compositional changes between dissimilar materials in layered composite structures.

By changing slurry compositions during freeze-spray process, one can achieve gradual compositional changes between dissimilar materials in layered composite structures. Such materials may find various applications particularly in high temperature environments where strong temperature gradients exist and require reduced internal stresses in graded composites. In this study, alumina-zirconia graded laminar composites are fabricated to demonstrate the feasibility of freeze-spray process followed by freeze-drying.Experimental procedureAlumina (A16SG: Alcoa, USA) and 5.3 weight% (3 mol%) yttria-stabilized zirconia(Sigma Aldrich, USA) powders were used for preparation of slurries. The average particle sizes of powders were d50: 0.40 µm for alumina and d50: 0.82 µm for zirconia.

The average particle sizes of powders were d50: 0.40 µm for alumina and d50: 0.82 µm for zirconia. Alumina and zirconia aqueous slurries with a solids loading of 30-50 volume% were prepared using 1.2 wt% ammonium polymethacrylate (Darvan C: Vanderbilt, USA) and 3 wt% acrylic emulsion polymer (Duramax: Rohm and Haas, USA). The amount of organic additives in weight percentage was calculated based on the weight of the powders. The optimum dispersant amount for slurries was determined by particle settling experiments. Well dispersed slurries were obtained by ball milling for 48 hours in polypropylene bottles. Viscosity of the slurries was measured using a rotating concentric cylinder viscometer (Haake, Model VT500) with shear rates at 0–500 s–1.The samples consisting of alumina and zirconia layers were obtained by spraying ofslurries on a cooled substrate. The thickness of each layer was controlled by repetitive spraying of slurries on frozen material.

The thickness of each layer was controlled by repetitive spraying of slurries on frozen material. Figure 1 shows a computerized 3D-gantry system equipped with a spray nozzle (Iwata, Model HP-BCS) used to spray thin layers of slurries on a cryogenically cooled metallic plate. The slurries freeze almost instantly, forming solid structures. The experimental set-up could be enclosed in a cooled chamber and dry remove the ice by sublimation. In order to study the effect of green density on the sintered microstructure, a set of samples were isostatically pressed at a pressure of 250MPa. After freeze-drying, the samples underwent a binder burn-out process at 550°C with a heating rate 1°C/min. Sintering of the samples was carried out at 1550°C for 6 hours with a heating and cooling rate 5°C/min.Figure 1. Experimental apparatus showing a system employed for sample fabrication processThe sintering density of the samples was measured by Archimedes method.

Experimental apparatus showing a system employed for sample fabrication processThe sintering density of the samples was measured by Archimedes method. Thermal expansion behavior of sintered materials was investigated using a dilatometer (Orton, Model 1600). Coefficient of thermal expansion (CTE) of materials was determined from the linear change in the length of the samples as a function of temperature. Microstructural and elemental analyses were conducted on the polished and fracture surfaces of the samples using a scanning electron microscope, SEM (JEOL, Model T330), and an energy dispersive x-ray spectrometer, EDS (Hitachi, Model S-4700 FESEM).Results and discussionFreeze-spray process of graded materials requires a delicate control of slurry propertiesand spray conditions. To achieve dense ceramic structures after sintering, the slurries shouldhave high solids loadings yet still have a low enough viscosity to be dispensable without causing clogging of the nozzle.

To achieve dense ceramic structures after sintering, the slurries shouldhave high solids loadings yet still have a low enough viscosity to be dispensable without causing clogging of the nozzle. Preparation of well dispersed and stable slurries is essential to prevent flocculation of the particles in the liquid which in turn affects the uniformity of the layer during deposition process and microstructural development of green and sintered bodies. Electrostatic, steric and electrosteric stabilization are common techniques employed in preparation of stable slurries17-19.Ammonium polymethacrylate was used as a dispersant to achieve steric stabilization inthe slurries. Alumina and zirconia slurries with solids loadings of 30, 35, 40, 45 and 50 vol% were prepared and characterized by viscosity measurements. Fig. 2 shows that all slurries exhibit shear thinning behavior.

Fig. 2 shows that all slurries exhibit shear thinning behavior. It is revealed that the increase of viscosity as a function of solids loading is more pronounced for zirconia slurries as compared to that for alumina slurries. Note that at high shear rates, the shear stress of the zirconia slurry with 50 vol% solids loading was too high for the sensor to be measured and not included in Fig. 2.(a)(b)Fig.2 Flow curve (viscosity vs. shear rate) of alumina (a) and ziconia (b)slurries showingthe effect of the increase in solid loading on the viscosityDispensing behavior of the slurries with the spray nozzle used in this study is stronglyaffected by the slurry viscosity which depends on the level of solids loading. Slurries with solids loadings up to 30 and 40 vol% could be sprayed uniformly without clogging of the spray nozzle. However, at higher solids loadings, slurries became highly viscous and were more difficult to spray continuously.

However, at higher solids loadings, slurries became highly viscous and were more difficult to spray continuously. As the solids loading of the slurries increased to 45 vol% and higher, clogging of spray nozzle was experienced, particularly for zirconia slurries. Based on the observations made from viscosity measurements and dispensing behavior of slurries, the optimum range of the solids loading was 35 to 40 vol% so that uniform spray pattern could be obtained without clogging of spray nozzle. The sintering density of alumina samples, fabricated using slurries with 35vol% solids loading, ranged from 75 to 86 % of theoretical density. To obtain higher densities, green samples were isostatically pressed at 250MPa. Isostatic compaction of freeze-dried alumina samples resulted in 90-98 % theoretical density after sintering. The microstructural development of sintered alumina prepared with and without isostatic compaction of freeze-dried samples is shown in Fig. 3.

The microstructural development of sintered alumina prepared with and without isostatic compaction of freeze-dried samples is shown in Fig. 3. It is evident that unpressed samples are highly porous while denser microstructures are obtained after isostatic compaction. Higher green and sintering densities could be obtained by spraying of slurries with higher solids loading and lower viscosity which could be prepared using organic additives.(a)(b)Fig.3 SEM micrograph of microstructure of alumina in an un-press (a) and isostatically-pressed (b) sampleLaminar alumina and zirconia composites were processed by freeze-spray deposition ofalternating layers as shown in Fig. 4. The EDS analysis of the samples revealed that the layers of dark contrast are alumina while the layers of bright contrast correspond to zirconia. Relatively pore-free alumina/zirconia interface is revealed.

Relatively pore-free alumina/zirconia interface is revealed. To obtain unique designs for microstructural development, laminated structures with smooth, wavy or more complex shape interfaces can be processed by adjusting the spray conditions during slurry deposition. Fig. 4 reveals that the thicknesses of alumina and zirconia layers are in the range of 200-300µm and 80-100µm, respectively. During deposition process, alumina and zirconia slurries were sprayed for 10 passes to constitute each observed layer. Each pass for spraying of slurries corresponds to formation of alumina and zirconia layers with a thickness of 20-30µm and 8-10µm, respectively. The layer thickness can be adjusted by controlling the solids loading of slurries, the flow rate and the speed of nozzle movement during spray process. (a)(b)Al2O3Al2O3Y-ZrO2Y-ZrO2Figures 4a and 4b. SEM micrographs of alumina/zirconia samples, showing uniform layer geometryFig.

(a)(b)Al2O3Al2O3Y-ZrO2Y-ZrO2Figures 4a and 4b. SEM micrographs of alumina/zirconia samples, showing uniform layer geometryFig. 5 shows the microstructure of a graded structure composed of alumina, alumina/zirconia composite, and zirconia layers. Layers A, B, and C correspond to alumina, alumina/zirconia (1:1 volume ratio, buffer layer), and zirconia, respectively. Such structures could be utilized in high temperature applications to reduce internal stress development within the material when composites are exposed to extreme temperature gradients at elevated temperatures20,21. As revealed in Fig 5, the interfaces between each layer are crack-free andappear to have a strong interfacial bonding. This can be attributed to alumina/zirconia buffer layer which forms a continuous alumina and zirconia at each interface.(a)(b)AAABBBCCFig.

This can be attributed to alumina/zirconia buffer layer which forms a continuous alumina and zirconia at each interface.(a)(b)AAABBBCCFig. 5 SEM micrographs showing a cross sectional view of a polished (a) and unpolished (b) alumina/buffer (50 vol% alumina + 50 vol% zirconia)/zirconia composite ACFigure 6. An alumina/zirconia composite (left) and an alumina/buffer/zirconia composite (right)Fig 6 shows that differential densification of alumina and zirconia layers during thesintering process may lead to warping of samples. Since densification of zirconia layer takes place at lower sintering temperatures as compared to that of alumina, excessive differential shrinkage due to constrained sintering may also result in crack formation perpendicular to the layers. The presence of a buffer layer can reduce the differential shrinkage so that less warping and crack-free samples can be obtained as shown in Fig 5.

The presence of a buffer layer can reduce the differential shrinkage so that less warping and crack-free samples can be obtained as shown in Fig 5. For the samples with planar geometries, the dimensional changes due to warping could be reduced by controlling of porosity of each layer in green samples or by applying pressure during the sintering process.To study the role of buffer composition on the bonding strength of interfaces, aluminaand zirconia bilayers with and without a buffer layer were prepared. It is known that macroscopic defects at the interfaces can cause de-lamination and result in catastrophic failure of the composites12. While the interface between zirconia and alumina layers in Fig 7 shows formation of voids or microcracking, the interfaces between alumina/bufferand zirconia/buffer layers are dense and free of micro-cracks (Fig. 8) as discussed previously.

10. The mean coefficient of thermal expansion was calculated using data collected between 250°C and 750°C. CTE values for alumina and zirconia were determined as 8.8×10-6/°C and 11.3×10-6/°C, respectively. It is revealed that the CTE increases linearly as the volume fraction (ranging from 0.2 to 0.8) of zirconia in alumina increases. Stress concentration at the interfaces caused by the mismatch of CTE between dissimilar materials can be reduced by the presence of a buffer layer which is composed of graded alumina/zirconia composites.Fig.10 Coefficient of thermal expansion of samples as a function of zirconia volume fractionSummaryA freeze-spray deposition technique coupled with freeze-drying process was developedfor rapid prototyping of alumina/zirconia lamellar structures.

Such components may find various applications particularly in high temperature environments where a wide range of temperature gradients is present and a reduction of internal stresses in composite materials is required.The process involves preparation of alumina and zirconia slurries with a solids loading in the range of 35-40vol% which were sprayed on a cooled substrate layer by layer to build desired structures. Rapid solidification of slurries with low binder content allows a rapid processing of composite materials. Freeze-dried parts were sintered and characterized by measurement of thermal expansion coefficients and studying of microstructural development. To achieve higher sintering densities, a set of freeze-dried samples were isostatically pressed.Alumina and zirconia bilayers were processed with and without a buffer layer which was prepared using composite powders of alumina and zirconia.

A graded interface between alumina and zirconia laminates using a buffer layer, was effective to reduce stress concentrations so that crack-free interfaces with strong bonding could be obtained. Linear dimensional change of the samples as a function of the temperature was determined by dilatometer measurements. Theresults were discussed with respect to microstructural development of graded ceramic composites.AcknowledgementThis work was supported by the Air Force Research Laboratories, Dayton, OH through the Center for Aerospace Manufacturing Technologies (CAMT) at the University of Missouri- Rolla. We also thank Drs. M. Leu, G. Hilmas, R. Landers, S. Reis at UMR and M. Hayes (Boeing) for useful discussions. D. Aiken, V. Satittavornchai, J. Mattingly, and C. Volek are acknowledged for their contributions to sample preparation.References1 M. Gasik, A. Kawasaki, and Y-S Kang, “Optimization of FGM TBC and Their ThermalCycling Stability,” Mater. Sci. Forum, 492, 9-14 (2005).

3Department of Mechanical Engineering – Engineering Mechanics, Michigan Technological University, Houghton, MI 49931. 4Department of Aerospace and Mechanical Engineering, University of Notre Dame, Notre Dame, IN 46556Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE- NA0002839.AbstractFielding Additively Manufactured (AM) parts requires evaluating both the part’s geometry and material state. This includes geometry that may be optically hidden. Both the geometry and material state affect the vibration response of the parts and modal analysis (identifying natural frequencies) has been shown to be effective for at least simple geometries using ex-situ methods (shaker table and impact hammer excitations).

This paper investigates evaluation of the frequency response of metal parts inside the build chamber using the process laser to excite the parts during printing (Renishaw AM250). Vibrations in the part are measured with accelerometers connected to the build plates and used to track the response during printing as during pauses between layers. The laser is modulated at different frequencies and focused onto specific targets to precisely extract the response from individual parts on the build plate. These results are compared to numerical models for metal parts of different geometries and with different material states. This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.14301.

IntroductionAdditive manufacturing is a growing means of producing parts due to its ability to create complex geometries that would otherwise be impossible to make with traditional methods. With the need for quality assurance being critical, evaluation of parts that are produced is an on-going field of study in the additive community. Evaluation schemes in the form of both destructive and non-destructive methodologies are the subject of research for both in and ex-situ analysis of these additively manufactured parts. Selective Laser Melting (SLM) is a common means of additively manufacturing metal parts; the range of materials that can be produced from SLM range from stainless steels, titanium, aluminum, nickel alloys, and composites.

With each material that is produced, one of the direst concerns is that of porosity, which can affect the quality and performance of the parts that are produced.The SLM process starts with a thin layer of powder that is spread over a build substrate at which point a focused laser scans over a pre-defined part geometry. The interaction between the laser and the layer of powder is a complex thermal process. The energy from the laser is absorbed into the layer via conduction and advection, resulting in fusion between the powder layer and substrate. The quality of this fusion is dependent on a key factor: recoil pressure; the origin of recoil pressure is due to the vaporization of the surface of the melt pool. With both the recoil pressure and Marangoni convection, the dynamics of the melt pool are described. Semek et al. reported that the vaporization-induced recoil pressure induces the ejection of the melt pool to form spatter [2].

Semek et al. reported that the vaporization-induced recoil pressure induces the ejection of the melt pool to form spatter [2]. [3] observed that the oscillations of recoil pressure resulted in the fluctuations observed in the formation of the rear keyhole wall. Using multi-laser technology, M. T. Andani et al was able to observe that with the increase in recoil pressure generated by multiple lasers, the amount of metallic material ejected from the melt pool is also increased [4]. [5] was able to experimentally observe the development in the melt pool using high-speed x-ray imaging and correlate it to the recoil pressure at that instant in time.In addition to experimental observations, various studies have been conducted in order to model and experimentally quantify the recoil pressure at the melt pool [6-8]. Simulating the interaction between the laser and powder bed involves incorporating both the thermal and dynamic interaction at the powder bed.

Simulating the interaction between the laser and powder bed involves incorporating both the thermal and dynamic interaction at the powder bed. Numerical simulations employ recoil pressure models that are based on the Clausius-Clapeyron model in order to better estimate how the recoil pressure influences metrics such as spatter ejection speed and temperature gradients at the melt pool [6,7]. As additional merit, [6] compared the ejection speeds of the particles in the simulation with experimental results using high-speed imaging and found that the behavior of both instances closely matched. By modeling the recoil pressure with this model, the behavior of the spatter and melt pool was able to be accurately described. C. Zhao et al. were able to observe changes in the spatter and correlated it to changes in the recoil pressure exerted on the melt pool [8].

C. Zhao et al. were able to observe changes in the spatter and correlated it to changes in the recoil pressure exerted on the melt pool [8]. This study estimated the recoil pressure in observation from literature by applying both the Clausius- Clapeyron model and assuming the pressure exerted onto the ejected spatter was directly related to the recoil pressure; the order of magnitude for the recoil pressure estimations was on the order of 60 kPa. This pressure can be made to be an equivalent force by multiplying by the average spot size diameter of the laser, in this case, 70 μm, which yields an equivalent force on the order1431of 0.25 mN.

The ability to measure the magnitude of the estimated recoil pressure is not difficult but rather the ability to measure the pressure over an area as small as the melt pool proves to be quite difficult.With recoil pressure being critical to the printing process, the ability to measure the pressure in- situ would allow real-time health monitoring of the melt pool. The ability to measure this pressure however, is difficult since the pressure is acting over a very small area. To calculate the recoil pressure, this study utilizes the frequency domain to calculation rather than temporal domain. The dynamics of the system given the force input from the laser shall be utilized from a modulated laser; the results from the measurements can then be extended to the Structural Health Monitoring (SHM) framework for NDT analysis of both the melt pool and parts that can be printed.14322. AM250 DynamicsThe Renishaw AM250 additive manufacturing system utilizes an Acousto-Optic Modulator (AOM).

AM250 DynamicsThe Renishaw AM250 additive manufacturing system utilizes an Acousto-Optic Modulator (AOM). This modulation allows for the control of the laser’s Pulse Repetition Frequency (PRF) through changes in the intensity of the sound wave that is diffracting the laser; this behavior is utilized for the purpose of being able to further control the melt pool. To understand the shape and periodicity of the waveform, a sample pulse mode was recorded using an experimental set- up involving a photodetector with a bandpass filter.

To understand the shape and periodicity of the waveform, a sample pulse mode was recorded using an experimental set- up involving a photodetector with a bandpass filter. The schematic of the experimental set-up, the same pulse mode, as well as the pulse mode in the frequency domain can be seen in Figure 1.Relative Angle between Laser and PDLaser Pulse8]V[e g a6t lSolid Angleo V4Melt Pool Area2OscilloscopeAlumina Disk300350400450 Time [(cid:541)s]500550600(a)(b)1]V[e g a t l o V0.10.019.09.510.0 Frequency [kHz]10.511.0(c)Figure 1: (a) Schematic of scattered light collection set-up (b) sample pulse modeFigure 1a shows the schematic in which the pulse mode was measured by incorporating a PDA100A2 photodetector and an oscilloscope. The radiance from the laser was collected by scattering it from an alumina disk.

The radiance from the laser was collected by scattering it from an alumina disk. Figure 1b is an example of a pulse mode for a given set of processing parameters and Figure 1c is the Fourier-Transform of the pulse mode. Modulating the laser allows for the laser to drive the part at a specific frequency and as a result, be able to record the response from the part and by extension, calculate the force and recoil pressure from the laser.14333.

Modeling Interaction Between Laser and Printed PartTo model the interaction between a part that has been printed and the laser, a simple Single Degree of Freedom (SDOF) mass-spring-damper model in which the forcing function was modeled as the laser, which can be seen below in Figure 2.Fx xTMBKFigure 2: SDOF model used to model interaction between laser and part.The equation of motion of Figure 1 is,+ (cid:6) Mx Bx Kx F+=(cid:6)(cid:6)(1)Where M, B, and K represent the mass, damping, and stiffness of the part and F is the forcing function due to the laser. For simplicity, the steady-state behavior of the system was found and transients were ignored.

For simplicity, the steady-state behavior of the system was found and transients were ignored. The canonical form of (1) was found by diving (1) by M which yields,( ) F t K ( ) F tω 2 n+ζω ω 2 (cid:6) x 2 n=+(2)(cid:6)(cid:6) xxntω=F sinoAssuming that the forcing function is of the form system is of the form ( ) x t system,and that the solution of the)(tω=sinX, (2) yields the steady-state form of the position of theX=K( 1−oF e )22 −rφ j+ζ 42 2 r(3)1434Where,φ⎧ ⎪ ⎪ = ⎨tan⎪ − π ⎪ ⎩ω ω n ⎞ ⎟ ⎠ ζ 2 r − 2 1 r=rζ 2 r − 2 1 r ⎛ ⎜ ⎝⎛ ⎜ ⎝− 1≤forr1⎞ ⎟ ⎠− 1tanfor r>1Plugging these values into the steady-state solution yields, ( ) − ω φ )22oF ( 1sin( ) (cid:6)(cid:6) x t=−Krt+ζ 42 2 rAt this point, the magnitude of the forcing function is required in order to model the dynamics of the system in steady-state.

The magnitude of the force was calculated by assuming that the energy from the laser was able to be converted into the kinetic energy of the melt pool, which can be seen below,P t laser laser on,eζ−t laser laser on,=1 2mmelt poolv2melt poolWhere P is the laser power, t is the period in which the laser is on, m is the approximate mass of the melt pool and (cid:535)(cid:3)(cid:76)(cid:86)(cid:3)(cid:87)(cid:75)(cid:72) damping ratio of the laser. The purpose of the damping ratio is to have a negative rate of change in the kinetic energy of the system.

The purpose of the damping ratio is to have a negative rate of change in the kinetic energy of the system. Taking the derivative of (5) with respect to the time that the laser is on and then multiplying by the mass of the mass pool yields the approximate force magnitude that would be exerted onto the melt pool,( F t olaser on,)=mmeltpoolP e laser−ζt laser laser on,( 12P t laser laser on,mmelt pool−et laser on,ζlaser−ζt laser laser on,)(6) is then plugged into (4), and then the double derivative with respect to time to calculate the acceleration of the part, yielding,( ) (cid:6)(cid:6) x t=K−ω 2 ( 1F o−rsin ) 22() − ω φt+ζ 2 4 partr2With (7), the ratio between the laser frequency and part frequency yields a maximum response when the part is driven near its natural frequency.

For NDT purposes, the response from the part when driven by a laser can be indicative of the recoil pressure as it is being printed.1435(4)(5)(6)(7)4. Experimental TheoryThe relationship between the FRF, acceleration, and force is,( FRF f)=( a f ( F f) )From (8), if the acceleration of a part as well as the FRF is known, the forcing function that would be exerted onto the part can be calculated. An example of an FRF can be seen with a simple tuning fork; an ANSYS simulation was conducted in which the tip of the prong was excited by a 1 N force and the acceleration in the same direction as the force was measured. The FRF from the tuning fork can be seen in Figure 3 below.0.100]0.080N m g [/0.060F R F0.0400.0202.55.010.0 Frequency [kHz]7.512.515.0Figure 3: FRF from tuning fork from ANSYS simulation.With Figure 3 the response of the part at resonance is several orders of magnitude higher than off-resonance.

With this tuning fork, the FRF is indicative of the part’s sensitivity to an input force at a range of frequencies. By leveraging the part near resonance, the acceleration that would be measured would be maximized.1436(8)5. Experimental Set-UpA build was created in which four sets of tuning forks were printed that were designed to resonate at a frequency: 4, 6, 8, and 10 kHz.

Experimental Set-UpA build was created in which four sets of tuning forks were printed that were designed to resonate at a frequency: 4, 6, 8, and 10 kHz. Three copies of each frequency tuning fork were printed, and they were arranged in rows, with Figure 4 showing a view of the build, with a number designation being seen in the figure to represent the frequencies of the tuning forks as well as the one that was excited being inscribed with a circle.Figure 4: Picture of tuning fork build for excitation experiment.Prior to being excited by the laser, an impact test was conducted on the 10 kHz tuning fork seen in Figure 4 for the purpose of seeing its first resonate frequency as well as having the FRF for the purpose of calculating the forcing spectrum; the FRF can be seen in Figure 5.

With respect to the coupling of energy between the part’s acceleration and the location of the accelerometer, an ANSYS simulation was ran in which the magnitude of acceleration was measured with a 1 N force input was measured at the same location as the force input and the same location as what is seen in Figure 4, with the ANSYS simulation results being seen in Figure 6.0.0250.020]N m / g [0.015F R F0.0100.0050.0002468Frequency [kHz]Figure 5: Experimental FRF of 10 kHz tuning fork.14371000100]g[n o10i ta r e e c c Al10.10.01Tip Build Plate0.001369Frequency [kHz]Figure 6: ANSYS FRF comparison when response from tuning fork was measured at tuning fork tip and build plate.From Figure 5, experimentally, FRFs are not as clean as a simulation would suggest. The first mode of the tuning fork that would later be excited occurred at 8.288 kHz, seen by the vertical peak in the graph.

The first mode of the tuning fork that would later be excited occurred at 8.288 kHz, seen by the vertical peak in the graph. With respect to the magnitude that would be measured by the accelerometer, an ANSYS simulation was conducted in which measurements were taken at the point of contact of the force and the location on the build plate where the accelerometer was placed for the experimental FRF. Figure 6 showcases the significant difference in magnitude at the first mode of the structure. For experimental purposes, placing the accelerometer on the build plate rather than the part itself serves two purposes: protecting the accelerometer from successive laser pulses and being able to monitor more than one part. For NDT testing, this placement of the accelerometer allows for continuous monitoring of parts as they are printing.For the excitation scheme, PRFs were selected based on being banded around the first mode of the tuning fork: 8.288 kHz.

Given that the resolution of the AM250 for PRF variance is on the order of 10 μs, the PRFs that were able to be generated were 7.1, 7.7, 8.3, 9.1, and 10 kHz. For each PRF pulse, data was collected from the accelerometer using a NI 9234 DAQ that sampled at 51.2 kHz.14386. Experimental ResultsThe acceleration spectrums that were measured for each PRF was compared against each other in Figure 7. The frequency band of interest was centered about resonance of the part with the closest PRF, 8.3 kHz, is seen as a vertical dashed line in the graph.10] g m[n o i t a r e e c c Al864PRF [kHz] 10 9.1 8.3 7.7 7.128.18.28.3 Frequency [kHz]8.48.5Figure 7: Measured acceleration spectrums for various PRFs.For the five PRFs, the maximum acceleration that was measured was 8 mg. The PRF that corresponded to this acceleration was the 8.3 kHz PRF. This is due to the laser’s frequency being close to that of the parts, resulting in the acceleration being maximized.

The PRF that corresponded to this acceleration was the 8.3 kHz PRF. This is due to the laser’s frequency being close to that of the parts, resulting in the acceleration being maximized. With each acceleration spectrum, Eq. 8 was used and the forcing spectrum was calculated by dividing acceleration spectrums by the experimental FRF of the tuning fork; the resulting forcing spectrums can be seen in Figure 8.1.6]N m[e c r o F1.41.21.00.80.6PRF [kHz] 10 9.1 8.3 7.7 7.10.40.20.08.18.28.3 Frequency [kHz]8.48.5Figure 8: Calculated forcing spectrums for various PRFs.1439The maximum force that was measured was 1.62 mN at resonance. Like the acceleration spectrum, this magnitude occurred when the PRF was closest to resonance. To relate the force measurement in Figure 8 to the recoil pressure, an average area of the spot size of the laser was used as the cross-sectional area of the applied force: 75 μm.

To relate the force measurement in Figure 8 to the recoil pressure, an average area of the spot size of the laser was used as the cross-sectional area of the applied force: 75 μm. The equivalent pressure was calculated to be 360 kPa. With the estimations from the literature that were calculated from this study being on the order of 60 kPa, the measured force and equivalent recoil pressure are on the same order of magnitude as estimations.7. Uncertainty in MeasurementsTo investigate the confidence in the magnitudes in both the force and recoil pressure, propagation of uncertainty was calculated. The uncertainty of the force was calculated first using the relationship between force, acceleration, and the FRF of the structure,σforce(f)=σ 2 laser 2 FRF( (f f) )+2 a laser) ( f FRFσ 2 FRF ( f4( )f)The uncertainty in the force was plotted with the frequency band centered around the resonance of the tuning fork below.

The summation of both terms in (9) was compared to just the first term based on looking into the difference in magnitude at resonance; both spectrums were plotted and can be seen in Figure 10 below for one of the PRFs as a means of behavior comparison rather than magnitude; the uncertainty spectrum can be seen in Figure 9.<JForce =10]N μ[1y tn ait r e c n U e c r o F0.10.010.0018.18.28.3 Frequency [kHz]8.48.5Figure 9: Uncertainty of force comparison at resonance between laser excitation and FRF components.At resonance, the uncertainty is proportional to the FRF of the structure: the magnitude of uncertainty is lowest at the part’s resonance. With this, the level of uncertainty can be minimized by driving the parts near resonance with the laser.

With this, the level of uncertainty can be minimized by driving the parts near resonance with the laser. The uncertainty spectrums were calculated for each PRF and were plotted about resonance and can be seen in Figure 10.1440(9)0.030]N m[y t n a t r e c n U e c r o Fi0.0250.0200.0150.0100.005PRF [kHz] 10. 9.1 8.3 7.7 7.10.0008.18.28.3 Frequency [kHz]8.48.5Figure 10: Force uncertainty spectrum for various PRFs.The magnitude of uncertainty when the PRF is closest to resonance is 31 μN. Since the magnitude of uncertainty scales with the FRF, the magnitude of uncertainty will always be largest with the PRF that is closest to resonance. The uncertainty in the recoil pressure was calculated by using a range of melt pool diameters from 50 to100 μm, yielding a standard deviation of 25 μm with the average melt pool diameter used be 75 μm. The propagation of uncertainty illustrated with Eq.

The propagation of uncertainty illustrated with Eq. 9, the uncertainty for the recoil pressure was calculated for the 8.3 kHz PRF,σRecoil Pressure=P Recoil⎛ ⎜ ⎝σforce F⎞ ⎟ ⎠2+⎛ ⎜ ⎝σarea A⎞ ⎟ ⎠2With Eq. 10, the uncertainty in the recoil pressure was calculated to be 41.3 kPa, leading to the recoil pressure at resonance for the closest PRF to be 360 ± 41.3 kPa. This recoil pressure is on the same order of magnitude as previous literature that dealt with quantifying and modeling the recoil pressure during the printing process.1441(10)8. Discussion and ConclusionsThe forcing function at resonance was calculated to be 1.62 ± 0.031 mN with the corresponding recoil pressure assuming variations in the melt pool diameter was calculated to be 360 ± 41.3 kPa. With literature estimations being on the order of 60 kPa, the equivalent force that would be expected to be measured is on the order of 1.1 mN.

With literature estimations being on the order of 60 kPa, the equivalent force that would be expected to be measured is on the order of 1.1 mN. The force and pressure calculations from this experiment are on the same order of magnitude as the estimations from literature. The absolute difference in magnitudes can be attributed to the structure that was used for excitation as well as the way the part was excited with the laser.With scope of this paper being NDT, a method for comparison is Acoustic Emission (AE). AE is a phenomenon that occurs when acoustic waves propagate through a medium when a solid undergoes a change in its internal structure; this can be seen to occur during the printing process as parts are formed at each layer. Studies have been able to observe changes in the acoustic signature of parts in both in and ex-situ circumstances [9-12]. L. W.

Studies have been able to observe changes in the acoustic signature of parts in both in and ex-situ circumstances [9-12]. L. W. Koester et al was able to differentiate between acoustic signatures with varying processing parameters in-situ during the SLM process [9]. The findings were able to be hypothesized to be able to correlate changes in the acoustic signature to defects that could potentially form during the printing process. [10-12] focused on ex-situ studies pertaining to the fatigue behavior of parts that were printed both additively and traditionally. [10,11] were able to observe a co-linear behavior between the acoustic signature measured and the growth of a crack in various types of metal plates. M. Strantza et al was able to observe an increase in the number of acoustic events, acoustic signals that go above a user-defined value, when the size of the crack maximized before failure.

M. Strantza et al was able to observe an increase in the number of acoustic events, acoustic signals that go above a user-defined value, when the size of the crack maximized before failure. With AE testing, the ability to monitor parts as they are printing can allow for diagnosis to be performed to determine if a part is fine or needs to be stopped due to a crack starting to form.Utilizing modal analysis for the purpose of a diagnostic method with laser excitation has good results about being able to quantify the recoil pressure. This is a simple extension of AE testing given that the structural response of a part will be measured during the printing process. It is not limited to measuring signals however, as it can be employed to the Structural Health Monitoring (SHM) framework by estimating the health of both the part and melt pool. The health of the part be estimations to the mass, stiffness, and damping matrices that describe the part layer-to-layer.

The health of the part be estimations to the mass, stiffness, and damping matrices that describe the part layer-to-layer. By being able to measure the forcing function of the laser and tracking the acceleration, the model matrices of the part can be calculated and compared to a baseline as well as correlating changes in the melt pool’s dynamics to transient changes in the magnitude of the recoil pressure.14429. AcknowledgementsThis work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1443References [1] M. J. Matthews et al., “Denudation of metal powder layers in laser powder bed fusion processes.” Acta Materialia. 114. 33-42. (2016). [2] V. Semak et al., “The role of recoil pressure in energy balance during laser materials processing.”, Journal of Physics D: Applied Physics. 30. 2541–2552. (1997). [3] M. J.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFrequency Inspection of Additively Manufactured Parts for Layer Defect IdentificationAimee Allen1, Kevin Johnson1, Jason Blough1, Andrew Barnard1, Troy Hartwig2, Ben Brown2, David Soine2, Tristan Cullom3, Douglas Bristow3, Robert Landers3, Edward Kinzel41Department of Mechanical Engineering-Engineering Mechanics, Michigan Technological University, Houghton, MI 499312Honeywell Kansas City National Security Campus, Kansas City, MO 641473Department of Mechanical Engineering and Aerospace, Missouri University of Science and Technology, Rolla, MO 654014Department of Aerospace and Mechanical Engineering, University of Notre Dame, Notre Dame, IN 46556AbstractAdditive manufactured (AM) parts are produced at low volume or with complex geometries.

Identifying internal defects is difficult as current testing techniques are not optimized for AM processes. The goal of this paper is to evaluate defects on multiple parts printed on the same build plate. The technique used was resonant frequency testing with the results verified through Finite Element Analysis. From these tests, it was found that the natural frequencies needed to detect the defects were higher than the excitation provided by a modal hammer. The deficiencies in this range led to the development of other excitation methods. Based on these results, traditional methods of resonant part inspection are not sufficient, but special methods can be developed for specific cases. This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1.

IntroductionAdditive manufactured parts have an advantage over traditional casting methods when considering parts produced at a low volume or parts with complex geometries. As the industry for additive manufacturing increases, new challenges become present. One of these challenges is part integrity. Defects on the surface of AM parts can be identified visually. Internal defects are more difficult to find as they cannot be seen by the naked eye.The AM process considered in this paper is selective laser melting (SLM). SLM is a 3-D printing process where a laser sinters each layer of metal powder to form 3-D printed metal parts. One internal defect that occurs from this process is a laser penetration defect or voids. Voids are holes in a part that occur in SLM when the laser misses sintering a layer or layers.The current methods for inspecting internal defects, such as voids, include computedtomography (CT) scanning and x-ray.

CT and x-ray testing are effective methods for detecting internal defects [1]. However, these methods take time to perform and can come at a high cost as1400CT and x-ray testing are not optimal for AM part inspection during and after processing [2]. Dynamic testing is a quick and low-cost method of testing for defects. The elimination of defects before CT and x-ray can save both time and money [3].Dynamic testing includes both shaker and impact (modal) hammer testing. AM partsexcited by these methods will vibrate at their natural frequencies. Natural frequencies, or resonances, are structural properties defined by mass, stiffness, damping, and boundary conditions for the structure [4].Near the resonance of the structure, the response amplitude will be higher than the otherfrequencies. The peaks on a Frequency Response Function (FRF) indicate the frequency at which a resonance occurs. An FRF relates the input to the output of two measurement points on a structure [4].

The peaks on a Frequency Response Function (FRF) indicate the frequency at which a resonance occurs. An FRF relates the input to the output of two measurement points on a structure [4]. The input for an FRF is usually a force with the output being displacement, velocity, or acceleration. One metric in which FRFs are often evaluated to determine if the measurement was “good” is by evaluating the coherence. The coherence is the amount of output directly related to the input. If all the input is directly related to the output the response is equal to one, if none of the input is related to the output the response is equal to zero.The FRF peaks can be represented by mode shapes. A mode shape is the physicalrepresentation of a natural frequency. For example, as suggested by Johnson et al., if a guitar string is plucked the 1st mode shape is the entire string vibrating and the 2nd mode shape is the string split in two with a node in the middle and the string moving on both sides of it.

Mode shapes become increasingly more complex as frequency increases. Due to this effect, the capability of finding tiny defects increases.Structures with the same material properties and boundary conditions will have the same natural frequencies and mode shapes. If the material properties or boundary conditions change, the peaks in the FRF will shift in frequency. A part containing a defect can be distinguished from good parts if its natural frequency is outside the standard deviation from a collection of good parts [3].Free-free boundary conditions are often the approach used to dynamically test structures. Free-free is suspending the structure on bungies or placing the structure on foam. By supporting the structure in free-free conditions, the modes from supporting structures, such as a fixture, are eliminated. Due to the increased cost and time of separating individual pieces for testing, the parts tested in this paper use a fixed boundary condition.Johnson et al.

Due to the increased cost and time of separating individual pieces for testing, the parts tested in this paper use a fixed boundary condition.Johnson et al. performed similar testing on chimney specimens. The authors concludedthat “dynamic evaluation method has significant potential to find ‘bad’ parts,” but further testing is needed before dynamic evaluation is commercialized.This article expands on the dynamic evaluation by Johnson et al. by considering a greaternumber of parts with increased complexity. Specifically, the research question investigated is “What size layer defects can be detected in 316L Stainless Steel parts attached to the same build plate?” The parts tested in this article were analyzed to determine if nominal groups of parts and1401groups with different amounts of defect could be separated from each other through resonant frequency testing.2. MethodsPart testing was the same as Johnson et al.

MethodsPart testing was the same as Johnson et al. except a small impact hammer and no Digital Image Correlation (DIC) were used. The small impact hammer was used as the parts were more closely spaced and smaller in size than the parts tested by Johnson et al. No digital image correlation was used since the initial shaker testing did not provide enough dynamic excitation nor distinct FRF peaks when recorded with a Scanning Laser Doppler Vibrometer (SLDV). The frequency range of interest for the parts tested in this paper were above the excitation range of the shaker. Additionally, Finite Element Analysis was used to validate the results from the experimental testing.The first set of parts tested was a build plate of 45 tensile bars provided by our research partner Missouri University of Science and Technology (MS&T). These tensile bars were split into nine groups of five. One group was nominal parts with the rest of the groups having a variation of the same layer defect.

These tensile bars were split into nine groups of five. One group was nominal parts with the rest of the groups having a variation of the same layer defect. This layer defect was a result of intentional missing laser sintering layers of powder. The shape of the defect was rectangular, representative of the geometry of the tensile bar (Fig. 1a). This layer defect ranged from 50 to 400 (cid:541)(cid:80) and was located approximately 1/3 of the way up from the bottom of each tensile bar (Fig. 1b). The defect increased in height for each subsequent group.Length = 0.21 inHeight = 4.7 inHeight = 1.75 inFigure 1: Height of tensile bar (1a) and defect length and height from base (1b)MTU was not informed which groups had each size defect until after testing andpresentation to the sponsor. The groups labeled 1-9 are shown from the top view of the build plate (Fig. 2a). The groups were randomly placed on the build plate to ensure no groups were congregated together.

The groups labeled 1-9 are shown from the top view of the build plate (Fig. 2a). The groups were randomly placed on the build plate to ensure no groups were congregated together. The bars were numbered for identification (Fig. 2b). These parts were not printed with the supporting structure. The supporting structure assists in printing the specimen and is removed from the build plate in post-processing.1402+v~ @ ,. ~ ,. ~ ,. ~ ,. ~~ ~ ~ ~ ,. ~ ~ ,. '- ~ ~ ,. ~ '- ,. ,. ~ ~ ,. ,.._ ~ ~ ,. ~ ~ ~ ~ '-@~,.._'- ~ ~ ~ '-,.._~,.._@'\lo,::,.,.. a,,'\,@@@(2a)(2b)Figure 2: Layer defect build with one group at nominal and the other eight with varied amount of layer defects provided by MS&TThe printed layer defect build tested is shown in (Fig. 3). Coupling between parts occurswhen parts are located near each other connected to the same structure.

3). Coupling between parts occurswhen parts are located near each other connected to the same structure. Weights were added to each of the bars not being impacted to eliminate the effect of coupling in the response. Each bar was impacted in the (cid:3398)X – direction with the SLDV recording the velocity response from the +X – direction. As only the x-direction of each cantilever beam was tested, the only modes that the results should display are x-axis bending modes. This means there should be no axial or y- axis bending modes in the experimental results. There is a possibility for some torsional modes to appear if the hammer impact was not directly in the center of the cantilever. The impact being off-center would cause some twisting in the structure producing a small torsional response.WeightFigure 3: Layer defect build with weights1403@a.,.

The impact being off-center would cause some twisting in the structure producing a small torsional response.WeightFigure 3: Layer defect build with weights1403@a.,. @The FEA modeling parameters consisted of modeling each individual bar on the plate.Each bar was fixed in the model where the bar connected to the build plate. Each bar was evaluated for up to 20 mode shapes.The next build plate provided to MTU was another forest tensile bar build with all thebars being nominal. This plate was tested in the same way as the first plate with the placement of each of the bars held the same. The purpose of testing this plate was to analyze the variation in the printing of the tensile bars and compare the results with the build plate with layer defects.The last build plate was three topology optimized brackets (Fig. 4). These parts wereprinted without defects in order to analyze whether the parts could be tested and if they compared to the FEA results performed on them.

4). These parts wereprinted without defects in order to analyze whether the parts could be tested and if they compared to the FEA results performed on them. If the parts could be tested and compared to the FEA, then defects would be added and analyzed. These parts were printed with the supporting structure since printing without the supporting structures failed. These specimens were tested in the ears at the top of the structure. The hammer direction and SLDV were oriented in the same direction as the forest bar builds.The FEA performed on the brackets was provided by Missouri S&T. The parameters for the FEA are similar to the forest bar builds. The boundary condition for the brackets was fixed to the plate. The first eight modes were calculated from the FEA.123Figure 4: Topology Optimized Brackets indicated by number provided by MS&T3.

The first eight modes were calculated from the FEA.123Figure 4: Topology Optimized Brackets indicated by number provided by MS&T3. Results1.1 Layer Defect BuildA resonant frequency test was performed to determine whether the 9 groups with defects could be distinguished from each other. These four modes, Mode 1, Mode 3, Mode 5, and Mode 8 were clearly distinguished from the resonant frequency test; whereas the other modes listed were determined from FEA (Fig. 5a). These four modes indicate the first four bending modes on1404the x-axis of the structure. There were no distinct bar groupings that could be visually identified (Fig. 5b and Fig. 5c). This indicates further testing is needed to determine if the groups can be distinguished.

There were no distinct bar groupings that could be visually identified (Fig. 5b and Fig. 5c). This indicates further testing is needed to determine if the groups can be distinguished. An interesting observation from this testing is that as the mode number increased there was a larger span between natural frequencies and more damping (wider peaks).100.00N/ ) s / N m / ( ) s e / m d u ( t i n g a Mg o L1.001.57N/ ) s / N m / ) ( s e / m d u ( t i n g a Mg o L0.01Legend: Group 1: Group 2: Group 3: Group 4: Group 5: Group 6: Group 7: Group 8: Group 9:20220320320420420520520620620769256950697070007020705070707093Hz Frequency (Hz)(5b)(5a)Hz Frequency (Hz)(5c)Figure 5: FRFs of the all bars on the defect build (5a), FRF at the 200 Hz mode (5b) and 7050 Hz mode (5c)The average and three-sigma for the bar groups from the first four bending modes wasperformed to establish if the defect groups could be separated.

The three-sigma for the nominal group on the build was compared to the other 8 groups with defects (Table 1). Since none of the other bar group’s natural frequencies (not shown) fell outside the three-sigma limits this indicates that no significant statistical difference was observed between the groups. This observation indicates that the natural frequencies of the bars from each group overlap other bar groupings and no individual defect groups were detected.Table 1: Three-sigma values for Group 1: Nominal parts on forest buildMode No.Natural Frequency (Hz)three-sigma (Hz)Mode 1205.35± 0.69Mode 31301.60± 14.53Mode 53565.28± 50.06Mode 87050.83± 85.193.2 FEA for defect buildThe experimental testing was verified through FEA on one bar from each defect group.The details on each of the groups were provided after the experimental testing results were shown to the project sponsor.

The defect size in each group was provided as follows: Group 1 has nominal parts (0 intentional defect) with each subsequent group increasing by a 50 (cid:541)m defect until Group 9 which has a 400 (cid:541)(cid:80) defect. From the experimental testing, the first four bending x-1405axis modes were found to correspond with modes in the FEA results (Table 2 not all modes are shown). This indicates that the experimental testing was valid as there was at most a 1.1% error between the results.The ninth mode was found to have the greatest difference between the nominal group andthe 50 and 400 (cid:541)m defects. This indicates that the first axial mode should be excited in order to differentiate the groups.Mode No.Frequency (Hz)Exp. Testing Avg.

This plate was tested using the same procedure as the layer defect build. The nominal bar build had more variation in natural frequency in comparison to the defect build (Table 3). The range and variance increased as the frequency increased for both builds indicating similar phenomenon in both plates. This indicates there is too much variation in the build process to determine the layer defect using the current testing method, since all the nominal bars should have the same natural frequency.Table 2: Range and variance for Group 1: Nominal bars compared to the same bars on the nominal build.Range (Hz)Variance (Hz2)Nominal buildGroup 1: NominalNominal BuildGroup 1: NominalMode 12.780.670.90.05Mode 317.874.8434.982.96Mode 541.0216.69187.0834.87Mode 889.9428.4883.68103.663.4 FEA for Topology Optimized BracketsThe FEA for the topology optimized brackets was provided to MTU before the bracketswere received.

The FEA was performed out to the 8th mode shape (Fig. 6). The largest motion for each of the modes appear to be in the ears of each of the brackets. This indicated that experimental testing should be performed on the ears. Another observation from the FEA results is that Bracket 1 and Bracket 3 appear to be similar as at least two of the mode shapes excite both1406of these brackets. This indicates that Bracket 1 and Bracket 3 will have some of the same natural frequencies in the FRFs.Mode 1: 5536 HzMode 2: 5716 HzMode 3: 6545 HzMode 4: 7272 Hz(6a)Mode 5: 7846 HzMode 6: 8082 HzMode 7: 8174 HzMode 8: 8324 Hz(6b) Figure 6: First 4 mode shapes (6a) and last 4 mode shapes (6b) for topology optimized brackets from FEA provided by MS&T3.5 Experimental Testing for Topology Optimized BracketsThe topology optimized brackets were tested in the same way as the nominal bar build todetermine the natural frequencies of these parts and compare them to the FEA results.

Only1407results from Bracket 1 are shown as the other two brackets were unable to be tested, as there was not enough excitation for the SLDV to be able to detect the response. Also, the small impact hammer was switched for the large impact hammer, as the small hammer did not provide enough excitation. Analyzing the FRF, there are no distinct peaks in the band of interest (5500-8500 Hz) for these brackets as there is noise introduced above 4500 Hz (Fig. 7). Also, the coherence becomes noisy and starts to trend downward around ~4500 Hz indicating that the energy inputted into the structure is not being measured in the response. These two observations indicate that the impact hammer did not provide enough excitation to the structure to determine the natural frequencies.

These two observations indicate that the impact hammer did not provide enough excitation to the structure to determine the natural frequencies. Therefore, there will be no comparison to the FEA from this testing.0.10NY1 Y2FRF Bracket:Point1:+X/Bracket:Point1:-X Coherence Bracket:Point1:+X/Bracket:Point1:-X/ ) s / m N( / ) e s d / m u ( t i n g a Mg o L1.00e-6 1.00/e d u t i lp m A0.00 010002000300040005000 HzFrequency (Hz)6000700080009000Figure 7: FRF and Coherence of Bracket 14. Discussion and Future Work4.1 Discussion of ResultsThis paper presented investigations into determining if defect groups in AM parts could be determined from resonant frequency testing. From both the experimental testing and the FEA results of the forest bar builds the defect groups were not able to be distinguished from each other.

From both the experimental testing and the FEA results of the forest bar builds the defect groups were not able to be distinguished from each other. Additionally, the topology optimized brackets were incomparable to the FEA since the resonant frequency test did not supply enough energy to the structure. A surprising finding was that the nominal build had more variation than the build with defects. In comparison to the study done by Johnson et al. the results in this study do not agree.The nominal build having more variation than the defect build was surprising. Onehypothesis on why this occurred is that unintentional defects were introduced in the printing1408process.

Onehypothesis on why this occurred is that unintentional defects were introduced in the printing1408process. A second hypothesis is that the bars had more coupling and interference with the surrounding bars when being tested than originally expected since, even with the weights, there was a high level of variation in the parts.The results from this study do not agree with the results from the chimney builds testedby Johnson et al., where they were able to determine which chimney had defects from a blind test. The difference in results between these two studies could be due to Johnson et al. investigating parts with much lower natural frequencies (less than 4000 Hz) with the parts being less complex and fewer in number per build in comparison to this study.

investigating parts with much lower natural frequencies (less than 4000 Hz) with the parts being less complex and fewer in number per build in comparison to this study. The results in this paper suggest that impact hammer excitation may be suitable for lower frequencies, but may not be suitable for parts with higher natural frequencies or increased complexity.4.2 Future WorkFuture testing for these parts will be aimed at providing more energy to both the forestplate builds and the topology optimized brackets. A testing technique that may be suitable on the forest plate builds is acoustic excitation in the axial direction, as this is the mode that will be most likely to distinguish the defects from one another. Additionally, ultrasonic testing may be suitable for testing these builds, but tends to be more costly and time consuming.

Additionally, ultrasonic testing may be suitable for testing these builds, but tends to be more costly and time consuming. The topology optimized brackets future testing will be focused on firing a high velocity projectile via a BB - gun at the ears in order to provide energy over 5500 Hz to the structure.5. ConclusionIn summary, the results from the investigations in this study were not able to determinethe defect groups in the forest bar builds nor provide enough excitation to the topology optimized brackets to be able to compare to the FEA results. These findings suggest that resonant frequency inspection may not be suitable for testing parts with high natural frequencies. Therefore, other excitation techniques need to be investigated in future testing. Future testing may involve acoustic excitation, ultrasonic excitation, or high velocity projectile testing.6. AcknowledgementsThis work was funded by Honeywell Federal Manufacturing & Technologies under Contract No.

AcknowledgementsThis work was funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.1409References[1] A. Karme, A. Kallonen, V.-P. Matilainen, H. Piili, and A. Salminen, “Possibilities of CT Scanning as Analysis Method in Laser Additive Manufacturing,” Physics Procedia, vol. 78, pp. 347–356, 2015.[2]NIST, “Measurement Science Roadmap for Metal-based Additive Manufacturing,” Energetics, 2013, https://www.nist.gov/sites/default/files/documents/el/isd/NISTAdd_Mf g_Report_FINAL-2.pdf[3]K. Johnson, J. Blough, and A.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFrequency Response Inspection of Additively Manufactured Parts for Defect IdentificationKevin Johnson1, Jason Blough1, Andrew Barnard1, Troy Hartwig2, Ben Brown2, David Soine2, Tristan Collum3, Edward Kinzel3, Douglas Bristow3, Robert Landers31Department of Mechanical Engineering-Engineering Mechanics, Michigan Technological University, Houghton MI 499312Kansas City National Security Campus, Kansas City MO 641473Department of Mechanical Engineering and Aerospace, Missouri University of Science and Technology, Rolla MO 65401Abstract: The goal of this paper is to evaluate internal defects in AM parts using dynamic measurements. The natural frequencies of AM parts can be identified by measuring the response of the part(s) to a dynamic input.

The natural frequencies of AM parts can be identified by measuring the response of the part(s) to a dynamic input. Different excitation methods such as a modal impact hammer or shakers can be used to excite the parts. Various methods exist to measure the parts’ responses and find the natural frequencies. This paper will investigate the use of Doppler lasers, accelerometers and Digital Image Correlation (DIC). The parts evaluated in this work include sets of parts that are still attached to the AM build plate, this makes the identification of a faulty part much more difficult as parts on a build plate interact with each other as well as the build plate complicating the responses. Several approaches to these issues will be presented based on the above listed response measurements.1. IntroductionAdditively manufactured parts (AM) have been shown to be a good choice for low volume uses that might not justify the cost of casting molds.

IntroductionAdditively manufactured parts (AM) have been shown to be a good choice for low volume uses that might not justify the cost of casting molds. Along with this increase in AM parts, comes new challenges. One of which is analysis of the AM parts for structural integrity. Defects on the outer surface of AM parts can easily be found with visual inspection just as they are with traditional subtractive manufacturing methods. AM parts can possess internal defects not typically seen in subtractive manufactured parts which also must be identified. Nondestructive test methods that are currently being used to find internal defects include x-ray and CT scanning. Although very effective, these methods can be very expensive and time consuming. A low cost method, with the potential to work for either low or high volumes of AM parts, is detection based on the dynamic properties of the parts. Although not as sensitive as X-ray or CT scanning, dynamic testing is inexpensive and fast.

Although not as sensitive as X-ray or CT scanning, dynamic testing is inexpensive and fast. Dynamic testing can be used as a prescreen for X-ray or CT scanning by quickly finding some defective parts which then do not need to be X-ray’d or CT’d as they are identified as “bad” parts prior to that step in the process, this has the potential to save both time and money.The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.2367If a build plate with AM parts is excited by means of a modal hammer or a shaker, the parts on the build plate will vibrate at their natural frequencies. Natural frequencies are defined by the ratio of stiffness to mass on a part as shown in Equation 1.𝜔𝜔 = √(𝑘𝑘/𝑚𝑚)the parts will vibrate with a displacement/velocity/acceleration amplitude much larger than at other frequencies.

These natural frequencies are indicated by peaks in the Frequency Response Function (FRF). A FRF is a function that relates the response of the structure at a particular location and direction to an input applied to the structure at a particular location and direction. The input is typically a measured force while the response is typically an acceleration or velocity. Anti-resonance frequencies are indicated by the valleys in the FRF shown in Figure 1. Assuming the input vibrational energy is evenly distributed, an anti-resonance frequency indicated little-to-no vibration at that frequency.Atthese frequencies, also called resonances,10000.001.00N / gg o Le d u t i lp m AFFRF Ch2-Loc5:+X/Ch2-Loc5:+X1.00e-30.000.00Hz4100.00Figure 1 – Example of a FRF of an AM part on a build plateIf identical parts have the same stiffness and mass, they will have the same resonant frequencies.

Respectively if identical parts have different stiffness or mass, some or all of the resonant frequencies will be different. Defective parts can be found by comparing the natural frequencies of each part with a standard deviation determined by a set of good parts.2368(1)For a natural frequency to change, a mass or stiffness variation in the part must be large enough and at a location which is significant for that natural frequency. This will be determined by the mode shape associated with the particular natural frequency, with some being sensitive to some locations of faults, and some not sensitive to the same faults. A mode shape is the physical shape at which the structure vibrates when the associated natural frequency is excited. For example, imagine strumming a guitar string.

A mode shape is the physical shape at which the structure vibrates when the associated natural frequency is excited. For example, imagine strumming a guitar string. The 1st resonant frequency mode shape is the length of the string vibrating back and forth, the 2nd mode shape is the string split in two with a node not moving and the string on each side vibrating back and forth. This increasing complexity in the mode shapes continues with higher and higher resonant frequencies. Only the resonant frequencies with the mass/stiffness variation involved in the mode shape at that frequency will be sensitive to that particular fault. Due to the increasing complexity of mode shapes as frequency increases, due to more localized portions of moving mass on the part, the ability to detect small defects improves at higher frequencies.The preferred method to test a part using this modal analysis method is with free-free boundary conditions.

Boundary conditions of this type would mean that the parts must first be separated from the build plate and then tested individually, this could add significant cost and time to the evaluation process. To save time and cost it is then advisable to test the parts while still attached to the build plate as multiple parts are tested simultaneously with this approach.2. Testing MethodsThe response methods used to measure the part response included several different technologies. The different technologies were chosen based on ease of use, quality of measurement, ability to discern small changes in mode shapes, and speed of testing. Accelerometers were used with both the impact hammer and the shaker input, these tend to be easy use and relatively low cost but are more time consuming in most scenarios if multiple locations are to be measured.

A Laser Doppler Vibrometer (LDV) was used to measure the response velocity on the parts while shaker testing, the laser had the ability to automate the scanning of responses at many different locations making it easy to measure a large number of different response locations. However, the LDV is an expensive piece of equipment. A pair of Digital Image Correlation (DIC) cameras measuring the response displacements of a grid pattern painted on the parts was also used. The DIC is the fastest method to evaluate a large number of measurement locations as all locations are measured simultaneously. The DIC used in this case had a frame rate of only 60 Hz and thus required a sinusoidal input to the shaker which is a slow method to excite many different frequencies to find natural frequencies.

The DIC used in this case had a frame rate of only 60 Hz and thus required a sinusoidal input to the shaker which is a slow method to excite many different frequencies to find natural frequencies. A set of high frame rate cameras is being investigated to speed this testing, with an obvious equipment cost penalty.The excitation methods used for exciting the AM builds were a modal impact hammer, which is an instrumented hammer that measures the force input to a structure or a large electrodynamic shaker as seen in Figure 2. The input signal to the shaker varied based on the technology used to measure the response of the parts. The shaker signals included periodic2369frequency sweeps for the accelerometers and LDV measurements and a sinusoidal input for the DIC measurement.cadbFigure 2 - Test Methods a) shaker, b) modal hammer, c) LDV, d) DIC3.

AM Build TestingAll build plates were provided to MTU as blind specimens without MTU being informedof what or where a defect was until after testing and presentation of results to the sponsor.3.1 Chimney 1The first AM build test consisted of five chimneys oriented as shown in Figure 3. One chimney was printed with a defect, it was unknown prior to testing what/where the defect(s) was. The chimneys were not printed solid to the build plate but rather with a support lattice structure, as shown in Figure 3. The support structure was developed to make it easy to remove the parts from the build plate after dynamic testing.237054321Figure 3 - 1st AM Build test – 5 chimneysThe Impact Hammer testing found the resonant frequencies of each chimney to be different by ~10 Hz and they were in order across the plate as numbered at some modes as shown in Figure 4.

It was determined that the AM printed parts had a slight difference in mass or stiffness due to their position on the build plate. At other modes, chimney #4’s resonant frequency shifted to a lower frequency, out of order with its position on the build plate.

This indicated a more significant property difference than that caused by the location on the build plate, as shown in Figure 4.1000.001.001000.00N / gg o Le d u t i lp m AN / gg o LF F F F FFRF Ch1-Loc5:+X/Ch1-Loc5:+X FRF Ch2-Loc5:+X/Ch2-Loc5:+X FRF Ch3-Loc5:+X/Ch3-Loc5:+X FRF Ch4-Loc5:+X/Ch4-Loc5:+X FRF Ch5-Loc5:+X/Ch5-Loc5:+XF F F F FFRF Ch1-Loc5:+X/Ch1-Loc1:+X FRF Ch2-Loc5:+X/Ch2-Loc1:+X FRF Ch3-Loc5:+X/Ch3-Loc1:+X FRF Ch4-Loc5:+X/Ch4-Loc1:+X FRF Ch5-Loc5:+X/Ch5-Loc1:+X10.000.001.001815.00Hz1965.002200.00Hz2400.00Figure 4 – FRF of Hammer Test Chimney Build 1 at 1900 Hz mode and 2300 Hz modeThe shaker testing agreed with the impact testing and also identified that the resonantfrequencies of chimney 4 shifted lower at some modes as shown in Figure 5.23711.00e d u t i lp m A0.00Figure 5 – FRF of Shaker Test Chimney Build 1 at 2300 Hz modeImpact FRF’s measured with a modal hammer are inherently less noisy than shaker FRF’s.

This can be seen in Figure 6. Therefore, it is desirable at times to use the modal hammer for finding the resonant frequencies. The difficulty with impact testing on AM parts still bolted to the base plate, is the modes of the base plate show up in the response of the parts on the build if the plate is tested in a free-free condition. This can be seen by the additional peaks (natural frequencies) in Figure 6 right hand graph that are not in the shaker data in Figure 6 left hand graph.1000.001.0010000.00g o Le d ut i lp m AN g/g o LFFRF Ch2-Loc5:+X/Ch2-Loc5:+XFRF Ch2-Loc5:+X/Sh-Table:+X0.10e-30.001.00e-30.00Hz4100.000.00HzFigure 6 – FRF of Shaker LH vs Hammer free-free RH from 0-4100 Hz.To acquire a clean FRF without the extra build plate modes, the build plate was bolted to the thick base plate of the shaker for an impact test.

This moved the base plate modes higher in frequency due to the added stiffness of the shaker baseplate and out of the frequency bandwidth of interest as shown in Figure 7.23724100.001.00e d ut i lp m A0.0010000.001.00N / gg o Le d u t i lp m AFRF Ch2-Loc5:+X/Ch2-Loc5:+X1.00e-30.000.00Hz4100.00Figure 7 – FRF of Hammer bolted to shaker bed plate.The LDV Test also showed that the natural frequencies of chimney 4 were lower at some modes. In Figure 8, the natural frequency of chimney 4 is approximately 50 Hz lower at the 2300 Hz mode, also shown is a still image of a video of build 1 shaking at chimney 4’s resonant frequency of 2289 Hz. Chimney 4 had significant amplitude in this picture while the other chimney’s exhibited much lower response amplitudes.Figure 8 – LDV data at 2300 Hz mode. Image from video at 2289 Hz.DIC was able to determine chimney 4 has resonant frequencies that are lower at some of the modes.

Image from video at 2289 Hz.DIC was able to determine chimney 4 has resonant frequencies that are lower at some of the modes. Since the cameras used are only capable of 60 fps maximum the test was not able to be done with a frequency sweep. The resonant frequencies were previously identified through an impact test for each mode, these frequencies were then used as shaker inputs to evaluate the mode shapes and response amplitudes. By exciting the plate at the natural frequency of the “good” parts, any parts not exhibiting large amplitudes would be the “bad” parts. Likewise, if it were excited, as in this case, at the natural frequency of a “bad” part the “good” parts would not exhibit large response. The still images from videos in Figure 9 were taken at 2288 Hz which is one of the resonant frequencies of chimney 4 as shown in Figure 8.2373aceFigure 9 – DIC at 2289 Hz.

The still images from videos in Figure 9 were taken at 2288 Hz which is one of the resonant frequencies of chimney 4 as shown in Figure 8.2373aceFigure 9 – DIC at 2289 Hz. a) Chimney 1, b) Chimney 2, c) Chimney 3, d) Chimney 4, e) Chimney 5The displacement of chimney 4 is high at this frequency and the mode shape replayed from the data visually more closely matches the true mode shape than do the other chimneys. The mode shape of chimney 1 is the next closest match as it is the next closest frequency to 2288 Hz.The chimneys on the first build were printed with support material as shown in Figure 5. During testing, the first resonant frequencies of the chimneys (the cantilever beam mode) changed between tests, this was identified to be a result of a fatigue weakening of the support material due to the vibration inputs used.

This result led to the realization that modes heavily influenced by2374bdboundary conditions may be of limited use in identifying parts with faults as the frequency could shift due to a boundary condition change and not a part fault.Build 1 conclusions include that chimney 4 was shown to have lower natural frequencies at some modes. This was due to chimney 4 being denser. Chimney 4 was printed with 150µm hatch while chimneys 1, 2, 3, & 5 were printed with 85µm hatch. The AM build printer laser power for all chimneys was 200W.3.2 Bracket BuildThe second AM Build consisted of five brackets shown in Figure 10. It was unknown how many had defects and what the defects were.

It was unknown how many had defects and what the defects were. Each bracket was printed with a support structure to the base plate.12345Figure 10 – Build with 5 brackets.The impact hammer and shaker testing found the natural frequencies of brackets 1, 3, and 4 were within 1 Hz of each other while brackets 2 and 5 were higher at some modes as shown at the 600 Hz and 2250 Hz modes in Figure 11.

In Figure 12, brackets 2 and 5 have a higher resonant frequency than brackets 1, 3, and 4.Figure 12 – LDV showing brackets 2 and 5 at higher resonant frequency at 600 Hz mode.Bracket build conclusions include brackets 2 and 5 were shown to have higher natural frequencies. This was due to being less dense. Brackets 2 and 5 were printed with 85µm hatch where brackets 1, 3, & 4 were printed with 150µm hatch all with 200W printer laser power.3.3 Chimney Build 2The second chimney build consisted of five chimneys printed solid to the base plate and inthe orientation shown in Figure 13. Again, it was unknown which chimneys had defects.23761.00e d ut i lp m A0.00Side 224135Side 1Figure 13 – 2nd chimney build.At the first bending mode, approximately 950Hz as shown in Figure 14, chimneys 1 & 2 are coupled together similar to a tuned-mass-damper.

Their resonance appear to be at the exact same frequency and they were moving 180 degrees out of phase, based on the mode shapes (not shown here). The stiffness and mass properties of chimneys 1 & 2 that affect this mode are the same. All of the chimneys also have sympathetic frequencies from the other chimneys’ natural frequencies. This is shown by the fact that each chimney’s FRF has a small peak at the natural frequencies of the other chimneys along with the large peak at their natural frequency. This makes the first bending mode difficult to use to find defects.

This makes the first bending mode difficult to use to find defects. In addition, chimney 3 seemed least affected by the others.1000.001.001000.00N / gg o Le d u t i lp m AN / gg o LF F F F FFRF Ch1-Loc4_5:+X/Ch1-Loc2_3:+X FRF Ch2-Loc4_5:+X/Ch2-Loc2_3:+X FRF Ch3-Loc4_5:+X/Ch3-Loc2_3:+X FRF Ch4-Loc4_5:+X/Ch4-Loc2_3:+X FRF Ch5-Loc4_5:+X/Ch5-Loc2_3:+XF F F F FFRF Ch1-Loc4_5_Side2:+X/Ch1-Loc2_3_Side2:+X FRF Ch2-Loc4_5_Side2:+X/Ch2-Loc2_3_Side2:+X FRF Ch3-Loc4_5_Side2:+X/Ch3-Loc2_3_Side2:+X FRF Ch4-Loc4_5_Side2:+X/Ch4-Loc2_3_Side2:+X FRF Ch5-Loc4_5_Side2:+X/Ch5-Loc2_3_Side2:+X1.000.001.00875.00Hz1035.00875.00Hz1035.00Figure 14 – 2nd chimney build side one left and side 2 right at 1st bending mode 900 Hz.At approximately 2000 Hz there is a breathing mode and torsion mode close to each other as shown in Figure 15. These two modes are the closest and interact the most on chimney 2.

Side 1 left and side 2 right. Image from LDV video bottom at chimney 4’s resonant frequency ~2890 Hz.In a comparison of Side 1 and Side 2 at 2000Hz as shown in Figure 17, it can be seen in the FRF’s taken on side 1 and side 2 of the chimneys that there are differences between the chimneys. Chimneys 1 and 4 are nearly identical side to side while chimneys 2 and 5 are not as similar.

a) Chimney 1, b) Chimney 2, c) Chimney 3, d) Chimney 4, e) Chimney 5,Chimney 2 build conclusions include the observance of sympathetic frequencies showing up on all chimneys indicating the first bending mode is not an ideal mode to use for finding defects. At the 2000 Hz mode there is a difference in chimney 2 that moved the breathing mode and torsion mode closer together than in the other chimneys. At the higher frequency modes, above 2800 Hz, chimney 2 starts to show significantly higher resonant frequencies relative to the other chimneys for a given mode. This frequency difference also seems to be greater side to side on chimney 2 than the other chimneys. The higher resonance frequencies in chimney 2 are due to it being built with 85 µm hatch while the other chimneys were 150 µm hatch. All with 200W printer laser power.2379bd4. Final ConclusionsFrequency response inspection can be used to find defective parts on an AM build.

All with 200W printer laser power.2379bd4. Final ConclusionsFrequency response inspection can be used to find defective parts on an AM build. Because of the support material failures and the sympathetic frequencies seen in some build plates, the lower frequency modes may not be ideal for finding faults. The higher frequency modes are not as affected by the support material or sympathetic frequencies of the other parts on the build and appear to be a better option for detecting defects. This is due to the mass of the chimneys at the build plate not being a significant part of the mode shape at some of the higher frequency modes. Since the higher modes have smaller local movements they are more sensitive to small stiffness/mass differences and therefore more useful in finding smaller defects.It was also shown that the boundary conditions of the base plate can have a large effect on the ability to easily identify part natural frequencies and therefore identify part defects.

Before using any of the excitation and/or response methods investigated a firm understanding of the part and build plate dynamics, natural frequencies and mode shapes, are necessary to successfully be able to identify part defects.Several different types of parts with different sizes and types of faults were evaluated inthis study. The range of parts/faults evaluated shows that this dynamic evaluation method has significant potential to find “bad” parts but further development is required to bring it to a viable commercial implementation.23804. References[1] R. J. Allemang. Vibrations; Experimental Modal Analysis. Structural Dynamics Research Laboratory, Department of Mechanical Engineering, University of Cincinnati, Ohio, USA, (2008)[2] R. J. Allemang. Vibrations; Analytical and Experimental Modal Analysis. Structural Dynamics Research Laboratory, Department of Mechanical Engineering, University of Cincinnati, Ohio, USA, (2008)[3]J. R. Blough, K. Johnson, A. Barnard.

el Cell Development using Additive Manufacturing Technologies - A ReviewN. P. Kulkarni, G. Tandra, F. W. Liou, T. E. Sparks, J. RuanDepartment of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65401 USAAbstractFuel cells are being perceived as the future clean energy source by many developed countries in the world. The key today for clean power is the reliance of fuel cells not only to power automobiles but also for residential, small commercial, backup power etc. which calls for production on a large scale. Additive manufacturing is perceived as a way to develop cost effective fuel cells. It imparts flexibility to design different kinds of fuel cells along with reduction in material wastage.

Additive manufacturing is perceived as a way to develop cost effective fuel cells. It imparts flexibility to design different kinds of fuel cells along with reduction in material wastage. This paper deals with the review of additive manufacturing processes for research and development of fuel cell components, such as synthesizing and prototyping new materials for fuel cell components, fuel cell system design and prototyping, designing well sealed fuel cells, bridging from fuel cell design to manufacturing tooling, etc.1. IntroductionFuel cells are electrochemical devices similar to batteries which convert energy from chemical state to electricity. There is an anode side and a cathode side in it. Fuel enters the cell from the anode side and oxidant flows into it from the cathode side. The reactants react inside the cell and the reaction products or the waste products flow out of it.

Fuel enters the cell from the anode side and oxidant flows into it from the cathode side. The reactants react inside the cell and the reaction products or the waste products flow out of it. The basic difference between batteries and a fuel cell is that fuel cell is only an energy conversion device and not energy storage device. Fuel cells consume reactant (fuel) from an external source which must be replenished. Hence fuel cells represent a thermodynamically open system. However batteries are both energy storage and conversion devices and hence they represent a thermodynamically closed system. The advantage of separating the storage and conversion functions is that power and energy capacity can be sized independently of each other.

The advantage of separating the storage and conversion functions is that power and energy capacity can be sized independently of each other. Also, many different fuels can be used as the primary energy source of the fuel cell setup depending on the types of fuels compatible with the type of fuel cell being employed [Spiegel 2006].This paper primarily discusses Proton Exchange Membrane (PEM) fuel cells because most of the research regarding fuel cells has been undertaken with regards to PEM fuel cells, due to its many advantages such as versatility. PEM fuel cells can be employed for various uses starting from portable power to automotive power to stationary residential power. The by-product of a PEM fuel cell is water, which is not only non- polluting but can be used as a potable water supply.In the PEM fuel cell, hydrogen is the fuel which enters the fuel cell through the anode end and oxygen through the cathode end.

The following reactions take place at the cathode and the anodeAnode: H2 (g)  2H + (aq) + 2e- Cathode: 1/2 O2 (g) + 2H+ (aq) + 2e-  H2O (l) Overall Reaction: H2 (g) + 1/2 O2 (g)  H2O (l)The components in a fuel cell are:1. Bipolar Plates. 2. Membrane Electrolyte Assemblies (MEA’s). 3. Gas diffusion electrode layers.Apart from this there are various auxiliary components such as gas flow pipes, the gaskets(seals), the connectors, end plates and cooling plates(required in fuel cell stack). The cost of these auxiliary components is relatively insignificant as compared to the cost of the major components.Table 1. Contribution of the components to the entire cost of the fuel cells [DOE 2005].Cell StackMembrane Catalyst Bipolar plates MEA’s35~40 % 15~20 % 10~15 % 30~35 %MEA is the heart of the fuel cell; rather it is the distinguishing criteria for different types of fuel cells.

An MEA, as the name suggests, is the assembly of the membrane and two electrodes on either side of the membrane. An electrode is a carbon cloth which is fabricated in a particular pattern depending on the mesh size required. Also, it needs to have specific properties to facilitate proper water management throughout the cell.As for the membrane, it is the electrolyte which is being employed for that particular fuel cell. The most common electrolyte used for a PEM fuel cell has been Nafion®. Nafion® is a generic brand name given by its developer DuPont. Its chemical name is sulfonated Polytetrafluoroethylene (PTFE). Although Nafion® is the most common polymer membrane employed in PEMFC, extensive research is being carried out to find a cost effective alternative which is as mechanically and chemically stable as Nafion® [Payne 2009].The most common catalyst used for the PEMFC is Platinum due to its stability in highly corrosive atmospheres as well as its performance characteristics.

The methods used for applying the catalyst are screen printing and hand painting. However, the uniformity of the catalyst deposited is not easily controlled. Also, these processes are time consuming, and require iterations of painting, drying and massing to achieve therequired loading of the catalyst. The reproducibility of these methods is poor. There is considerable amount of catalyst wasted in the feed lines due to clogging which results in an increase in the production cost [Taylor 2007].The aforementioned commercial methods of producing major components of fuel cells are not in accordance with the economic threshold value as required by the US Department of Energy.

These processes combine costly materials and processes that result in increased costs of fabrication of fuel cells.To achieve the target of production cost of $30/kw by 2015 as set by the US DOE [DOE 2005], there is a need to achieve low cost fabrication of fuel cells and use alternate cheaper materials in the manufacturing processes. Based on the Results of the Workshop on Manufacturing R&D for the Hydrogen Economy, several challenges confront the transformation of the U.S. manufacturing sector to support the hydrogen energy economy such as:Develop innovative, low-cost manufacturing technologies for new materials and material applications.Adapt laboratory fabrication methods to low-cost, high-volume production.Rapid manufacturing is an innovative manufacturing technique which can be used for the fabrication of fuel cells which goes hand-in-hand with the aim of US DOE.

Rapid prototyping is defined as a machine technology which is used to fabricate 3-dimensional models and prototype parts from a numerical description (typically a CAD model) using an additive approach to form physical models. That is why Rapid prototyping is also referred to as ‘Additive Manufacturing’. Additive Manufacturing (AM), as the name suggests is the process of fabrication of physical models or prototypes by addition of materials. This addition takes place layer by layer incrementally. By this process, the problems of form generation and material composition are addressed. The smaller the incremental volume of material better is the accuracy of the form generated and also the control over system parameters. AM doesn’t require any external tooling for the manufacturing of 3D freeform objects.There are various kinds of widely used Additive Manufacturing techniques suchas:1. Selective Laser Sintering(SLS) 2. Fused Deposition Modeling (FDM) 3. Stereolithography (SLA) 4.

Selective Laser Sintering(SLS) 2. Fused Deposition Modeling (FDM) 3. Stereolithography (SLA) 4. 3D printing 5. Laminated object manufacturing (LOM) 6. Electron Beam Melting (EBM)Variations of these processes also exist but it is not important to be listed above since a small variation of some system parameters might lead to an entirely different process. For the manufacturing of PEM fuel cells, more importance has been given to a few processes such as Fused Deposition Modeling (FDM), 3D Printing and Selective Laser Sintering (SLS) which are discussed in detail in this paper.The advantages of AM itself make it an attractive way to build fuel cells. With additive manufacturing technologies, you have the flexibility to change the design of the fuel cells without the need to change the entire setup as would be required with regards to conventional manufacturing technologies. In this paper, there is an example of a planar array fuel cell with a mono polar plate design.

In this paper, there is an example of a planar array fuel cell with a mono polar plate design. It gives a good proof of the flexibility of additive manufacturing technology. This feature of AM enhances the prospects of further cost reduction. Inkjet printing aids the process of precision manufacturing since we can deposit materials with micrometer precision thereby again reducing material waste. Impressive results from the three processes as described in this paper maximize the scope of AM for building fuel cells. It might happen that, under a single roof, we see multiple AM techniques used to build an entire fuel cell.2.

It might happen that, under a single roof, we see multiple AM techniques used to build an entire fuel cell.2. FDM (Fused Deposition Modeling)Fused Deposition Modeling is an additive fabrication technology which constructs superior rapid prototypes from 3D CAD data where in a thermo plastic material is extruded in the form of beads layer by layer using a temperature controlled head which is actually controlled by Computer Aided Manufacturing (CAM) software [Zhong 2000].A plastic filament or wire supplies material to an extrusion nozzle which is heated so as to melt the material and deposit the required amount of material in horizontal and vertical directions(i.e., wherever it is necessary).The material hardens as soon as it is extruded from the nozzle.The thermo plastic materials used in FDM process have good stability and durability of the mechanical properties over time; they have high heat resistance and also produce smooth parts with all the finest details intact.

The commonly used materials with this process are Acrylonitrile Butadiene Styrene (ABS) polymer, elastomers, investment casting wax and some of the water soluble materials are used in this process which acts as support structures during the manufacturing process [Masood 2004].Fig. 1. Process of FDMFDM has been used for the fabrication of miniature fuel cell stack in a planar array form [Chen 2008]. For the development of miniature fuel cells, it is required to have pin-point precision, since the aim of this type of fuel cell is to have high power density in a small stack. A study has been made to develop a 10 cell planar array air breathing fuel cell using FDM as the RP (Rapid Prototyping) process [Chen 2008].Fig.2. Construction and layout of the stack [Chen 2008].The alternative processes for the development of such miniature fuel cell are the Micro-Electro-Mechanical processes (MEMS) and the conventional CNC machining processes.

The flow field plates were the parts which were fabricated using RP, more specifically Fused Deposition Modeling (FDM). Acrylonitrile-butadiene-styrene (ABS) was the material used for the fabrication of flow field plates since it has high mechanical strength, low cost and easy to fabricate by RP.The following table gives a comparison of the manufacturing time required for the geometry of flow field plates by different processes [Chen 2008].Table 2. Comparison of different methods for fabricating the flow field geometry [Chen 2008].Process Rapid Prototyping CNC MEM’sTime (approx) 1 hour 2hours 12-36 hoursIt is noted that RP is faster than the rest of the methods. Also, if the flow field plate is designed more and more complicated, CNC machining may not be possible at all.

Also, if the flow field plate is designed more and more complicated, CNC machining may not be possible at all. This kind of miniature PEM planar array FC stack is a first try in both academic as well as industrial world.The FDM process used for planar array fuel cell fabrication is described next withrespect to the design and performance of the PEMFC.2.1 Design and characteristics of the planar array PEMFCThere are 10-segments in the PEM. Hence the total reactive area is 17cm2 temperature  The cell temperature is 25oC.Two configurations have been tried: Parallel and series.  These 2 configurations have been tested with natural convection and forced convection.The anode is on the same side of the membrane whereas the cathode is on the opposite end or the ventilated end.Hence, it is called as a mono-polar stack design.2.2 Performance test in parallel connection [Chen 2008].Fig.3. Comparison of forced and natural convection in a parallel connection.

Comparison of forced and natural convection in a parallel connection. [Chen 2008]2.3 Performance test in series connection [Chen 2008].Fig.4. Comparison of forced and natural convection in a series connection. [Chen 2008]ambientParallel connection stack has higher power density than serial connected stack since some cells performing badly will affect serial connection where parallel connection won’t be affected a lot. [Chen 2008]. The performance of the stack reaches power density of the state of the art planar array fuel cells (100-120 mW/cm2) [Chen 2008].Clearly, fuel cell components made by RP (FDM) instead of conventional CNC machining or more costly MEM processes do deliver performances as required which does speak about the reliability of the process. So, we infer that RP is a successful procedure in prototyping the components. In the future, we might even see RP being applied for larger scale production.3.

So, we infer that RP is a successful procedure in prototyping the components. In the future, we might even see RP being applied for larger scale production.3. 3D Inkjet printing3D Inkjet printing is yet another form of Additive Manufacturing. What differentiates 3D printing from other forms of additive manufacturing is that it is much more affordable than other processes existing till date. Inkjet printers are plug and play devices that require little setup, training or maintenance.Inkjet printing utilizes drop-on-demand technology to deposit various materials in a colloidal ink form. Also, there is no contact between the printer head and the substrate on which it is going to be deposited. There are two types of inkjet printers- one which use piezoelectric transducers and one which use thermal resistors to expel droplets through the nozzles.

There are two types of inkjet printers- one which use piezoelectric transducers and one which use thermal resistors to expel droplets through the nozzles. Development of inkjet printers will result in smaller nozzle sizes and hence ink droplets, which will result in higher resolution (dots per inch) as well as in printing intricate features, patterns which is advantageous in the development of fuel cell components.Inkjet printing can be employed in printing different MEA’s since the composition is not very different from each other. Inkjet printing can be considered as an efficient method used for the deposition of catalyst layers because of the performance it gives in terms of controlled catalyst deposition for ultra loadings of Platinum which results in a better utilization of Pt as compared to conventional catalyst deposition methods like Screen Printing and Hand Painting. Inkjet printing will also help in optimizing the Pt loading which will result in reduction of costs.

Inkjet printing will also help in optimizing the Pt loading which will result in reduction of costs. The reproducibility produced in the catalyst printing is incredible and this will in turn lead to lesser cell failure rates.Fig.5. Sample catalyst layers printed using inkjet printing [Taylor 2007]3.1 Inkjet printing as compared to other printing techniques [Taylor 2007].IJP as compared to other printing techniques proves to be more advantageous as it allows for a uniform distribution of catalyst material onto the surface of GDL and provides picolitre precision and control of the deposition of each print and thus paves a way for ultra low loadings. IJP is also found to be reproducible due to the elimination of some of the intermediate steps of drying and massing which are two important steps in Hand painting and screen printing.Fig.6.

IJP is also found to be reproducible due to the elimination of some of the intermediate steps of drying and massing which are two important steps in Hand painting and screen printing.Fig.6. Time illustration of inkjet printing compared to hand painting [Taylor 2007].Catalyst inks should be similar to the OEM inks as specified by the manufacturer so that the printing can be executed smoothly.Table 7. The usual properties for home/office printer inks [Towne 2007].Viscosity1-4(centipoises) Surface tension (mN/m) Average Particle size (μm)30-35 0.23.2 Setup used for the experiment [Towne 2007].Catalyst formation takes place by thoroughly mixing a carbon supported catalyst with Nafion® solution and de-ionized water. Water, ethylene glycol and isopropanol are added to achieve the required properties of surface tension and viscosity.

Water, ethylene glycol and isopropanol are added to achieve the required properties of surface tension and viscosity. The Nafion® membrane is prepared by washing in 3% wt H2O2 for 1 hour, rinsed and boiled in de- ionized water for 1 hour and stored in Milli-Q grade de-ionized water. The printers considered for this experiment were simply off the shelf printers whose cartridges were cleaned off the original ink and replaced with the catalyst ink with the help of a syringe. Illustrator software is used for making different size and shape electrodes and for different amounts of platinum loading by changing the hue, saturation and luminescence.3.3 Results and discussion [Towne 2007].Table 8. The ink characteristics [Towne 2007].Property Particle sizeSurface Tension ViscosityResult < 2µm35.5 mN/m 3.35 cPWithin range NoNo YesComments Not within range but the ink is maintaining colloidal stability.

Just a tad out of range.There were two attempts to print the catalyst layers. One was on a cellulose acetate substrate and the other was on Nafion® substrate.Fig.7. SEM cross section of a single catalyst layer on cellulose acetate substrate [Towne 2007].Fig.8.SEM cross section of eight printed catalyst layers on Nafion® [Towne 2007].In the above the catalyst layer is seen between the substrate and its thickness is 1.02 μm thick since below the thickness of 580 the catalyst nm, the is increases detrimental to our interests.the epoxylayer andresistivity of tremendously whichIn the above figure a layer thickness upto 3.2 µm was measured. We can observe here that the thickness of the catalyst layer is not uniform and is very uneven. This happens due to swelling of the Nafion® membrane.As expected, the Nafion® substrate swells due to the water and alcohol in the catalyst ink. Water alone can lead to swelling of the membrane by 32%. This results in uneven printing.

Water alone can lead to swelling of the membrane by 32%. This results in uneven printing. The cross section of the single layer catalyst section as seen in figure b clearly shows the swelling of the membrane.Fig.9. Optical micrographs taken at 15x magnification showing evidence of banding in three samples ofdifferent thickness [Towne 2007].The three figures above show 3 single layer inkjet printed catalysts. They have different amounts of thicknesses (drop amounts). The darker the layer, the thicker it is and hence well connected. This leads to better conductivity.Thus it is evident that inkjet printing allows excellent control over the individuallayer thicknesses. Hence, many layers and ultimately thicker electrodes can be deposited.The earlier images show that it is difficult to print the catalyst layer on a Nafion® membrane, however, with some post processing, the catalyst layer can be made uniform as well as well mechanically adhered to the membrane.

The usual post processing steps are hot pressing and water extraction.Hot pressing leads to removal of ethylene glycol and also it leads to more uniform catalyst formation on the membrane.Fig.10. TEM image of printed catalyst layers on Nafion® before processing [Towne 2007].Fig.11. TEM image of printed catalyst layers on Nafion® after hot pressing [Towne 2007].The arrow represents carbon particles. This figure depicts the discrete nature of platinum and carbon particles. This definitely affects the interconnectivity and thus the conductivity.You cannot make out the separate layers of carbon and platinum particles, thus showing the continuity. Hot pressing was done at 2045 psi, 125oC for 5 min3.4 Performance of single cells [Towne 2007].The testing of these catalyst printed membranes was carried out by making asingle cell out of it. MEA’s had printed layers on both anode and cathode.Table 9.

MEA’s had printed layers on both anode and cathode.Table 9. Specifications of the MEA [Towne 2007].Active area of the MEA Platinum loading Drop size2.25 cm2 0.094 mg/cm2 3 pLOnly 2.75% H2 was used for initial studies for comparison of different electrodes. The testing started by the comparison of the unprocessed MEA with hot pressed MEA’s.Fig.12. Graph comparing printed MEA with processed MEA’s [Towne 2007].It was presumed that hot pressing did not remove the ethylene glycol completely. Hence water soaking was carried out on the printed MEA to remove the rest of the ethylene glycol. Water soaking led to maximum current output of 106mA/cm2 at 0.401V. Thus the density of power is 42.4mW/cm2.After this initial testing, 100% H2 was used for better comparison with the commercial MEA’s. The figure below describes the comparison of different MEA’s with 100% H2 and with different treatments after the printing.Fig.13.

The figure below describes the comparison of different MEA’s with 100% H2 and with different treatments after the printing.Fig.13. Power Curves comparing a commercial MEA with printed MEA’s [Towne 2007].In figure 13, it is clear the commercial MEA’s outperformed those inkjet printed MEA’s are having the catalyst loading of 0.094 mg Pt/cm2. However, when the catalyst loading was 0.2 mg Pt/cm2, it was a comparable performance as compared to commercial MEA’s.Table 10. Comparison of the improved MEA with the commercial MEA [Towne 2007]. Platinum loading ( mg Pt cm-2 ) 0.2 0.3Type of MEAPeak power density ( mWcm-2 ) 155 167Inkjet printed CommercialHence, with a 33% lower catalyst loading, only 7% lower power density was obtained. This result proves that inkjet printed MEA’s can compete with the commercial ones.Thus, it is evident how efficient can inkjet printing method for fabricating MEA’s is as compared to commercial MEA’s.

The efficiency of the catalyst usage or loading can be further enhanced by grading the amount of platinum loading in every layer. Previous literatures suggest that the graded catalysts were found to perform better than the uniformly loaded catalyst in every layer [Xie 2005, Wang 2004].Paganin et al. clearly suggests that platinum is better utilized when it is more concentrated near either the GDE layer or the electrolyte membrane layer [Paganin 1996]. Inkjet printing makes it possible to grade the platinum loading print after print. Previous research carried out by Taylor et al. demonstrated that a graded catalyst of Pt wt% 10-50 on carbon black outperformed the uniform catalyst structure of 20 % wt Pt on carbon black at nearly the same amount of overall platinum loading [Taylor 2007].Fig.14. Graded Catalyst layer [Taylor 2007]. uniform catalyst to a graded catalyst [Taylor 07].Fig.15. Performance comparison of a standard4.

Graded Catalyst layer [Taylor 2007]. uniform catalyst to a graded catalyst [Taylor 07].Fig.15. Performance comparison of a standard4. SELECTIVE LASER SINTERINGThe functions of bipolar plate are:To provide electrical contact between two adjacent MEA’s.  Uniformly distribute hydrogen gas and oxygen gas/air to the anode and the cathode side of the MEA respectively.To serve as a platform to support the soft MEA.  To act as an outlet medium for heat and water vapor generated from the net reaction.Hence, the requirements for a bipolar plate accordingly are:High electrical conductivity  Plate material electrically compatible with the electrode  Very low gas permeability for reactant gases.  High thermal conductivity to make use of the waste heat.  Chemical stability i.e. corrosion resistant.  Low density plate material to keep the stack volume and weight low.

 High thermal conductivity to make use of the waste heat.  Chemical stability i.e. corrosion resistant.  Low density plate material to keep the stack volume and weight low.  Inexpensive plate material.4.1 Material selection for a fuel cell bipolar plateThere are three types of materials identified for the manufacture of fuel cell bipolar plates which are: pure graphite, metallic materials, and carbon-polymer composites.Pure graphitePure graphite is a very good conductor of electricity with peak conductivity of 1.44x103 S/cm. Graphite is very difficult to machine when it comes to the machining of the flow field channels because of its flaky microstructure and irregular geometry. This also reduces its mechanical strength [Chen 2006].Metallic materialsMetal powders such as stainless steel, titanium, gold, aluminum, have good machining characteristics as compared to graphite. However, gold and titanium are very costly. Aluminum can be used with a gold coating.

However, gold and titanium are very costly. Aluminum can be used with a gold coating. However, there is large difference in co-efficient of thermal expansion which leads to micro-cracks in the coating. Stainless steel has corrosion issues [Maeda 2004, Chen 2006].Composite materialsComposite materials suitable for the application of bipolar plates are a combination of porous graphite along with polycarbonate plastic. Graphite is an allotrope of carbon and a semimetal. The carbon based materials suitable are resins such as polyethylene, phenolic, Vinyl ester etc. with filler materials like carbon black and carbon/graphite powders.

The carbon based materials suitable are resins such as polyethylene, phenolic, Vinyl ester etc. with filler materials like carbon black and carbon/graphite powders. These composite systems provide electrical conductivity as well as corrosion resistance and mechanical strength [Chen 2006].4.2 Set-up and procedure for the experiment.There are two kinds of SLS procedures: Direct and IndirectDirect SLS means parts are produced by just laser sintering of the powder withoutany post processing measures Indirect SLS involved production of a porous green part held together by a certain polymer binder followed by some post processing measures.According to the research conducted by Chen, indirect SLS of carbon based composite material accommodates the material and procedure selection criteria for the PEM fuel cell bipolar plate fabrication [Chen 2006]. This Indirect SLS proceeded in 3 stages to meet all the plate requirements:1. SLS of bipolar plates 2.

This Indirect SLS proceeded in 3 stages to meet all the plate requirements:1. SLS of bipolar plates 2. Carbonization of the binder 3. Epoxy infiltrationTable 11. Key process parameters for SLS process [Chen 2006].Powder constituentsComposition: Average particle size Powder mixing SLS machine CO2 laser power Laser scan speed Laser scan spacing Powder layer thickness Powder bed preheating Temp Purging gasGraphite (GrafTech GS150E) and Phenolic resin (Georgia Pacific GP5546) 70w% graphite and 30 w% phenolic resin Graphite: 80 μm Phenolic resin:11μm Roller mixed DTM Sinterstation 2000 10~20 W 60 in/s 0.003 in 0.004 in 60˚C NitrogenAfter the SLS process, post processing of the bipolar plates was further carried out. The post processing basically consisted of two steps of binder carbonization and epoxy infiltration.CARBONIZATION PROCESS [Chen 2006].A vacuum furnace was used for this purpose.

The post processing basically consisted of two steps of binder carbonization and epoxy infiltration.CARBONIZATION PROCESS [Chen 2006].A vacuum furnace was used for this purpose. The maximum heating capacity of the vacuum furnace being 2000˚C.Argon gas was filled into the chamber to prevent oxidation of carbonized phenolic resins which reduce the glassy carbon yields.Table 12. Temperature rise and the ramp rates [Chen 2006]. TEMPERATURE Room Temp-200˚C 200˚C-600˚CINITIAL PROFILE INTERMEDIATE PROFILE FINAL PROFILEAt 800˚CRAMP RATE 60˚C/hr 30˚C/hr0At 800˚C, the dwelling time was 1 hour. During this process the phenolic binders are burned off and a part of it was converted into glassy carbon. This resulted in good interconnected pores which increased the electrical conductivity as well as the mechanical strength of the parts.EPOXY INFILTRATION FOR FINAL SEALING [Chen 2006].After the carbonization process, the structure was still found to be porous.

Epoxy resin was chosen as the infiltrant to seal these pores because of its good mechanical strength, chemical stability and ability to wet most substrate materials. Clear coat resin which is a mixture of more than 70% diglycidyl ether of bisphenol A and less than 30% alkylglycidyl ether was used for this purpose. The resin was initially cured with the help of a hardener and then diluted with solvents like toluene, xylene etc., this was done in order to reduce the viscosity of the resin so that it can easily penetrate through the cured pore structure. The epoxy, hardener and the solvent should be mixed in proper ratios (2:1:1) to avoid the formation of un-reacted epoxy and hardener which affect the final So as to form a gas tight plate structure the brown part were part properties. immersed in the infiltrant at least twice. The parts were then oven dried at 60˚C for several hours to remove residual moisture.

immersed in the infiltrant at least twice. The parts were then oven dried at 60˚C for several hours to remove residual moisture. The electrical conductivity of the infiltrated parts was found to be better than the brown parts.4.3 Results and discussionTable 13. Properties of the SLS bipolar plates [Chen 2006]. TEST METHODPROPERTYVALUEFlexural strength Electrical Conductivity Specific weightCorrosion Rate Gas PermeabilityTwo point bend test Four point probe test Archimedes principle of fluid displacement Tafel extrapolation method leak spectrometer Mass detector1730psi 80 S/cm Avg. density=1.27g/cm36µA/Cm2 5x10-6 Cm3/Cm2.sInterfacial contact resistance<200mΩ.cm2 / 1.6MPaAll the above properties were found to be satisfactory but the electrical conductivity of these bipolar plates could be improvised to meet the target set by the DOE [DOE 2005].

The following methods were followed to improve the electrical conductivity [Chen 2006]: Addition of a liquid phenolic infiltration/re-curing step prior to final sealing  Reduction of glassy carbon resistivity by curing process parameter controlInfiltration of brown parts with conductive polymerThese processes showed results which are quantified as belowTable 14. Enhancement in electrical conductivity.First Infiltration/Recurring stepSecond Infiltration/Recurring step~108 S/Cm (35% boost in the conductivity) ~117 S/Cm (8.3% further enhancement)Fig.16. Graph showing the improvement in electrical conductivity with each process [Chen 2006].5. ConclusionsThe paper reviews three Additive Manufacturing (AM) processes. Each process is suitable for building specific fuel cell components. Performance characteristics of these components fabricated using AM processes prove that they give performance equal or better than the components fabricated by conventional techniques.

Performance characteristics of these components fabricated using AM processes prove that they give performance equal or better than the components fabricated by conventional techniques. Inkjet printing, amongst all AM processes is the process where most of the research has been undertaken with regards to building fuel cell components. The reason for that is it is easiest to commercialize as compared to rest of the methods since off the shelf printers have been demonstrated to produce components competitive with the commercial ones. FDM is convenient for planar array fuel cells as compared to MEMS and CNC manufacturing. Indirect SLS of graphite composite bipolar plate demonstrates fabrication of plates having superior characteristics which also meet DOE specifications. Development of these processes on a commercial basis is still under a lot of investigation.6. ACKNOWLEDGEMENTSThis research was supported by a grant from the U.S.

Development of these processes on a commercial basis is still under a lot of investigation.6. ACKNOWLEDGEMENTSThis research was supported by a grant from the U.S. Air Force Research Laboratory contract #FA4819-09-C-0018. Support from the Missouri S&T Intelligent Systems Center is also greatly appreciated.7.

LL-DENSIFICATION OF SLS PARTS BY RE-MELTINGT. NIINO and H. YAMADA Institute of Industrial Science, The University of Tokyo 4-6-1 Komaba Meguro Tokyo, 153-8505 JapanReviewed, accepted August 4, 2004AbstractAmong commercially available rapid prototyping processes, SLS is the most effective in terms of adaptability of various materials. However, rapid prototyped parts by the process are always porous and the physical properties of the parts are different from dense parts which is to be used in final product. This paper introduces a post process that can densify SLS processed plastic parts to almost 100%. An SLS processed polystyrene part is densified and, resultantly, a much stronger and transparent part is obtained.IntroductionSince market debut of stereolithography, various types of rapid prototyping processes and devices have been developed, and some are still being.

In the early period of RP’s history, the most effort was focused on how to build parts successfully without failure, and then how to make them accurately. However, recent discussions have shifted to how to realize good qualities other than accuracy such as tensile strength, heat resistance, chemical toughness and possibility of building out of special materials like metals or ceramics.In terms of part build accuracy, stereolithography is superior to selective laser sintering (SLS). On the other hand, in many cases, SLS is more convenient because of its faster building speed and applicability to larger variety of materials such as high-temperature polyamide and easy-to-be-lost polystyrene.

While SLS’s good ability of building prototypes out of various plastics provides good estimation of chemical property of their product version, which is to be injection-molded in the most cases, prediction of physical or mechanical performances such as tensile strength is poorer because of inevitable porosity of SLS processed parts. For example, an SLS part out of CastForm PS are much more fragile than a typical injection-molded polystyrene part, and handling of the part requires great care if it is not reinforced by wax infiltration. From a viewpoint of estimating a prototype in advance of making die and mold, the prototype should be fully dense and made out of the same material as its product version.400Investment casting of plastics using plaster mold is a classical way of obtaining fully dense parts. However, it is not always easy to burn out the model especially when the structure of model is complicated.

However, it is not always easy to burn out the model especially when the structure of model is complicated. The authors propose a novel process to densify an SLS parts by wrapping it with plaster and then remelting.This paper introduces the process for full-densifying SLS processed parts and reports process parameters of experimental densification of polystyrene parts. Some physical parameters such as tensile strength of densified parts are measured and reported.ProcessFig. 1 shows outline of the process that can densify SLS processed parts. First, an SLS processed part is invested with plaster. After hardening the plaster, the investment including the SLS part is loaded into a vacuum chamber that is installed in an oven. The vacuum chamber is evacuated first, and then the oven’s heaters are turned on to melt the parts.

The vacuum chamber is evacuated first, and then the oven’s heaters are turned on to melt the parts. Doing these in this order allows the air out of the pore network in the SLS part to be removed before the molten resin out of the part blocks the network to leave isolated pore containing the air in the middle of the part; otherwise, bubbles would remain in the final densified part. In this process, the molten resin flows into the lower section of the plaster cavity and remain there, while, in investment casting process, it should flow out of the cavity instead. Since the original SLS parts are porous, the whole cavity cannot be filled up only with original model. To obtain desired shape of full-dense parts, a dummy volume out of the same material should be attached at upper part of the desired shape of the SLS part.

To obtain desired shape of full-dense parts, a dummy volume out of the same material should be attached at upper part of the desired shape of the SLS part. The dummy can be built as a part of SLS model.Dummy volumePour plasterMolten resin flowHeatingFlaskEvacuationPut model in a flaskInvestmentPut investment in vacuum chamberCool & Breaka.Build SLS modelb.Investc.Remelt model of resin in vacuumd.Obtain dense partFig. 1 Densification ProcessExperimentsHeating and evacuating systemFig. 2 depicts composition of heaters and vacuum system. The vacuum chamber is made from stainless steel and all the gasket is oxygen-free copper. Thus the chamber resists against high temperature above 300°C. Capacity of the chamber is approximately 9400cm3 (9.4l), and it is evacuated with a rotary pump of which evacuation rate is 1700cm3/s (100l/minute). The401chamber is equipped with two systems of thermocouple feedthroughs.

The401chamber is equipped with two systems of thermocouple feedthroughs. The chamber is placed in an oven whose highest temperature is 270°C.Heating temperature and timeOvenDensification of parts out of CastForm PS was experimentally performed. To test the process, test pieces as shown in Fig. 3 were fabricated with Vanguard HS. The test piece consists of two sections, a cube and a sphere. The two sections are connected with a short cylinder, and the cube and the sphere are used as the desired shape and the dummy volume, respectively.Vacuum ChamberThermocouple180270that affect successful densification are heating temperature and time for which the model isThe mostimportant parametersVacuum PumpFig. 2 Composition of Heating and VacuumSystemkept at the temperature. Higher temperature gives better liquidity to the molten resin, and longer heating improves replication of the shape.

2 Composition of Heating and VacuumSystemkept at the temperature. Higher temperature gives better liquidity to the molten resin, and longer heating improves replication of the shape. On the other hand, both of them encourage deterioration of the resin, and excessive temperature or heating time yellows the densified parts. Thus the temperature and time must be minimized.05 85 1To find the lowest temperature that can liquidify the resin and make it flow into the lower cavity, preliminary test was performed. One gram of the powder including small lump as shown in Fig. 4a is put in a oven, and time when powder of CastForm is measured. At 180°C, the powder melted and the surface becomes flat in 15minutes as shown in Fig. 4b. Contrarily, in case of 170°C, some lumps are still remaining even after 1hour, and the surface of the powder cannot become flat(Fig. 4c). Therefore, we defined the lowest temperature that liquidifys CastForm as 180°C.0 250[mm]Fig.

4c). Therefore, we defined the lowest temperature that liquidifys CastForm as 180°C.0 250[mm]Fig. 3 Dimensions of the model used indensification test402a) Original powderb) after15 minutes at 180°Cc) after 1 hour at 170°CFig. 4 Result of Preliminary Test to Find the Lowest Liquidifying TemperatureTo find optimal heating time, four types of heating profiles (type A through D) were tested.(cid:122) Type AAfter evacuating the chamber that was containing the investment, the thermostat of the oven was liquidifying the temperature of 180°C. Fig. 5 displays variation of temperatures of the air in the oven, investment’s surface and resin. As shown here, it took 10 hours for resin to reach 180°C. After that, another 3 hours was spent for full densification.set atlowest]C o[e r u t a r e p m e T210180150120906030005 Time [hours]Resin Mold Air1015As a result of this densification, the densityFig.

5 Variation of temperature at each point in type A experiment.of the parts increased from 0.371g/cm3 to 1.051g/cm3 reaching typical value of polystyrene. In the obtained part, no hollow whose diameter is larger than 1mm was observed. However, microscopically small bubbles are remained. (cid:122) Type BA higher temperature of 270°C, which is the maximum of the oven, was tested. As shown in Fig. 6, it took also 10 hours to reach 270°C. After reaching 270°C, heater was turned off and the chamber was cooled down with the oven’s door opened. Until the temperature of the resin downed to 180°C, inside of the chamber was kept vacuum the resin from deterioration or to prevent yellowing. After 1.5 hours, the temperature of the resin lowered to 180°C. Then, the chamber was vented and the investment was picked out.

After 1.5 hours, the temperature of the resin lowered to 180°C. Then, the chamber was vented and the investment was picked out. In this experiment, full densification had completed without keeping resin temperature at 270°C for a certain period because the resin’s temperature have been higher than 180°C for more than 6 hours. (cid:122) Type C300240]C o[180e r u t a r e p m e T120Resin Mold Air600 0510 Time [hours]15Fig. 6 Variation of temperature at each point in type B experiment.To minimize the required time for reaching 180°C, oven’s temperature was set at the403highest of 270°C until the resin temperature reached 180°C. After it reached, the vacuum chamber’s surface was cooled down to 180°C rapidly by turning off the heaters and opening the oven’s door to keep the resin temperature at 180°C. Then, thermostat was set at 180°C. As a result, the resin temperature was controlled as shown in Fig. 7.

Then, thermostat was set at 180°C. As a result, the resin temperature was controlled as shown in Fig. 7. Afterward, by keeping resin temperature at 180°C for 4 hours, the part was fully densified, compressing the total densifying time to 9 hours without heating up the resin above 180°C. (cid:122) Type D240]C o[180e r u t a r e p m e TResin Air Mold1206000510 Time [hours]15Fig. 7 Variation of temperature at each point in type C experiment.To reduce the total processing time more,following procedure was followed: 1. Start heating with the thermostat of the ovenset at its maximum of 270°C2. Heat the resin up to an adequate temperature]C o[e r u t a r e p m e T24018012060Resin Mold Airabove 180°C03. Turn off the heater and open the oven’s doorto cool the resin down to 180°C4. Pick the investment out of the vacuum051015Time [hours:Fig.

Turn off the heater and open the oven’s doorto cool the resin down to 180°C4. Pick the investment out of the vacuum051015Time [hours:Fig. 8 Variation of temperature at each point in type D experiment.chamber and break it when the temperature comes to 180°C By this procedure, the total process time was successfully reduced by 2.5 hours reaching 6.5 hours. Variation of the temperature in this experiment is displayed in Fig. 8. In this experiment resin temperature reached 212°C at maximum.Deformation during densificationDifference in length at each point of the model as shown in Fig. 9 was measured as displayed in Tbl. I. In the densifying the previous section, two types of original pre-densified parts were used. The difference between these two types is only date when the models were built. There was no intentional difference in building parameters.tests describedinAs for length l and depth d, deformation caused by densification was very small as 1% or less.

There was no intentional difference in building parameters.tests describedinAs for length l and depth d, deformation caused by densification was very small as 1% or less. It is even404Tbl. I Deformation of each point after densificationProcessCAD Design Original part 1 Original part 2Type AType BType CType DOriginal Part Data TypeN/A N/A N/A1222Length [mm] Length [mm] Length [mm] Length [mm] Deviation [mm] Error Rate [%] Length [mm] Deviation [mm] Error Rate [%] Length [mm] Deviation [mm] Error Rate [%] Length [mm] Deviation [mm] Error Rate [%]l 50 50.6 50.4 50.6 0.0 0.0 52.0 1.6 3.2 50.6 0.2 0.4 50.6 0.2 0.4Point d 20 20.6 20.1 20.5 -0.1 -0.5 22.1 2.0 10.0 20.3 0.2 1.0 20.3 0.2 1.0h 15 15.2 15.1 14.3 -0.9 -5.9 15.8 0.7 4.6 14.5 -0.6 -4.3 14.7 -0.4 -2.6smaller than that between CAD design and SLS parts. At the most point, length was increased. This seems to be caused by grains of plaster stuck on the parts. As shown in Fig.

At the most point, length was increased. This seems to be caused by grains of plaster stuck on the parts. As shown in Fig. 10, a layer of plaster grain was stuck on the surface of the model. This plaster cannot be removed by plaster solvent, which is used in typical investment casting. Amount of stuck plaster tends to increase when process temperature is high.hWith regard to height h, large decrease was observed. Shrinkage encouraged by absence of holding pressure seems to cause this decrease.dlFig. 9 Measured length1mma) Plaster stuck on a model. (Partially removed)b) Microscope photo of stuck plasterFig. 10 Plaster Stuck on a Densified PartTensile strengthTbl. II Mechanical Property of Densified partstest was carried out on densified parts processed under various conditions, and their tensile strengths and Young’s modulus were measured as listed in Tbl. II.

II. As shown here, shorter process time at lower temperature prevents deterioration of the resin giving larger tensile strength and Young’s modulus. The maximum strength of 13.5MPa was one third or quarter of that of typical injection-molded polystyrene.Tensile ruptureDensification ConditionTensile Young's Strength Modulus 759MPa 13.5MPa 300MPa 9.4MPa 188MPa 6.8MPaTemp. 180℃ 4 hours 180℃ 15 hours 270℃ 15 hoursTimeLight transmittance and colorFig. 11 shows densified parts that are processed at 180°C for 11, 7 and 4 hours, respectively. After the process, each part was knocked out of its investment and machined into a cube of 10mmä10mmä45mm. As a result of full-densification, the parts became transparent since boundaries of refractive index, there used to be a lot of which between CastForm PS and the air,405through were eliminated.

As a result of full-densification, the parts became transparent since boundaries of refractive index, there used to be a lot of which between CastForm PS and the air,405through were eliminated. Light paths of 5mm at wave length of 540nm were 63%, 53% and 59% respectively for the process times of 11, 7 and 4 hours.transmittancesab.The part that had been kept melted for 11 hours, model (a) in Fig. 11, appeared more yellow than the others to naked eye. However, there was no significant difference in light transmittance measurements at five wave lengths of 420, 460, 510, 540 and 600nm.c.Fig. 11 Transparence of densified parts. Process times for parts a, b and c are 11, 7 and 4 hours, respectively.Remaining bubblesAs mentioned in the previous section, no hollow was observed in densified parts. However, small bubbles, diameters of which could be measured as around 40µm were found as shown in Fig. 12.

However, small bubbles, diameters of which could be measured as around 40µm were found as shown in Fig. 12. Each bubble seems to be containing thermal decomposition gas.200µmFig. 12 Bubbles in the partDiscussionAs a problem, it took longer time for the investment loaded in the chamber to be heated up to the melt temperature than we expected. This is because of strong heat insulation of vacuum. In order to shorten this duration, heating system can be revised. For example, putting some heaters in the vacuum chamber and placing them near the investment or attaching them on the surface of the investment.Since polystyrene is heated at high temperature for long period, the resin was deteriorated even if the circumstance was maintained vacuum. This deterioration weakened and yellowed the resin, and made small bubbles reducing transparency of the parts. To avoid this, the time keeping the resin at high temperature should be reduced.

This deterioration weakened and yellowed the resin, and made small bubbles reducing transparency of the parts. To avoid this, the time keeping the resin at high temperature should be reduced. Pressurizing with inert gas such as argon after the resin is melted and pore network was blocked can be expected to enhance the replication speed. This can also reduce deformation caused by shrinkage.Sticking plaster is very troublesome especially in densification of fine and complicated structure since finishing of which is difficult. There seems to be no easy solution, but using RTV as release agent was successful.406ConclusionsA process that can increase density of SLS processed porous parts to the material’s original density remaining the shape of the model was introduced. The process was tested on densification of SLS parts out of polystyrene. Resultantly, it is found that the resin must be heated to 180°C at lowest and the temperature must be maintained for 4 hours at least.

Resultantly, it is found that the resin must be heated to 180°C at lowest and the temperature must be maintained for 4 hours at least. If the temperature is higher, the maintaining duration can be reduced. By this process, the density of the model was increased by a factor of 2.8. With regard to deformation during the densification, it is very small as 1% or less in horizontal direction. On the other hand, that in vertical direction is very large as 3 to 4% because of shrinkage. Tensile strength of the model was dramatically improved, but it did not reach that of material itself because of deterioration caused by high temperature during densification. Densification gave transparency, but the model was hazier and more yellow than injection molded polystyrene. The densified model did not include hollow inside but micro bubbles of which diameters are around 40µm.407.

LLY RECYCLABLE MULTI-MATERIAL PRINTING Jonathan D. Hiller, Hod Lipson Mechanical and Aerospace Engineering Cornell University Ithaca NY, 14853, USAAbstractRecycling is often a costly and inefficient process, particularly for objects composed of multiple integrated materials. Here, we demonstrate a freeform fabrication system that prints with fully reusable physical voxels and minimal recycling effort. This new paradigm of digital (discrete) matter enables any number of materials to be printed together in any configuration. The individual voxels may then be reclaimed at will by dissolving the bonds holding the structure together. Coupled with a compatible voxel sorting process, we demonstrate multiple generations of freeform fabricated objects using the same physical material.

Coupled with a compatible voxel sorting process, we demonstrate multiple generations of freeform fabricated objects using the same physical material. This opens the door to a flexible desktop fabrication process in which 3D multi-material objects are fully recyclable and re-usable with minimal infrastructure.IntroductionWith the advent of multi-material additive manufacturing (AM) processes and the recent push for sustainability among developed nations, the need for recycling of additively fabricated 3D parts has become a critical need for future research (Bourell et al., 2009). The AM industry is growing steadily (Wohlers, 2008), but as AM parts see widespread adoption the potential impact has not been thoroughly considered (Yanchun et al., 1999).

The AM industry is growing steadily (Wohlers, 2008), but as AM parts see widespread adoption the potential impact has not been thoroughly considered (Yanchun et al., 1999). Here, we focus specifically on recycling end-use additively manufactured objects, not the full lifecycle analysis of additive manufactured parts (as in Jansen and Krause, 1995, Drizo and Pegna, 2006 and Hopkinson et al., 2006).The vast majority of parts created with additive manufacturing processes today are not recycled. This is primarily for two reasons. First, in many AM processes the material in use undergoes irreversible changes such as polymerization or infiltration with binder which makes further re-use of the material intractably inefficient. Second, AM parts made of material which could theoretically be recycled, such as those using thermoplastic or metals, are not recycled due to lack of volume, incentives and infrastructure.

Second, AM parts made of material which could theoretically be recycled, such as those using thermoplastic or metals, are not recycled due to lack of volume, incentives and infrastructure. To recycle these parts would require an infrastructure involving an energy- intensive melting stage, along with specialized equipment to return the material to its original form, such as extruded thermoplastic for fused deposition modeling (FDM) or powder of an appropriate size distribution for laser sintering. Although technically feasible, this is not common in practice, mainly due to the low volume of AM parts and the lack of economic incentive.The challenges of recycling multi-material objects extend far beyond the realm of additive manufacturing. Vast amounts of electronics waste are laboriously and hazardously decomposed by hand in developing countries (Hilty, 2005).

Vast amounts of electronics waste are laboriously and hazardously decomposed by hand in developing countries (Hilty, 2005). More generally,many manufactured parts consisting primarily of recyclable material cannot be recycled because separating the materials is often prohibitively expensive (Sodhi and Knight, 1998). A manufacturing process which addresses these challenges could have significant impact on how waste is handled. Additionally, remote situations such as space exploration would benefit from completely recyclable materials (Malone and Lipson, 2006, Robert et al., 2002).The results presented here should not be confused with the process of recycling material involved in an AM process but unused in the final part. For instance, it is common practice to reuse the leftover resin in a stereolithography build, and it is also possible to recycle the unused bed of powder in power-based processes (Pham and Gault, 1998).

The challenges in re-using and recycling these unused materials are wholly different than those needed to recycle the material used to make the actual end-use part as we address here.BackgroundDigital materials are a new paradigm in which 3D matter is composed of elemental physical building blocks (voxels), as illustrated in Figure 1 (Hiller and Lipson, 2009a). In the same way as a digital signal is represented as a series of logical ones and zeros, a 3D object is represented completely by the presence or absence of a voxel at each location within its 3D lattice. The advantages of the digital realm observed in other technologies such as computing and communications have direct parallels in the world of 3D matter.

The advantages of the digital realm observed in other technologies such as computing and communications have direct parallels in the world of 3D matter. An object is perfectly repeatable over an infinite number of generations and error correction can be employed to create "perfect" objects, but at the expense of increased processing effort and a finite resolution.(a)(b)Figure 1: Traditional analog materials (a) are continuous in nature and every dimension has an associated finite error. Digital (discrete) materials are composed of fundamental, aligned building units (voxels) and the structure is defined perfectly by the presence or absence of a voxel at each location within the lattice.When this abstract idea of digital materials becomes concrete as part of a physical voxel-based manufacturing process, a number of additional advantages are also realized.

(Hiller and Lipson, 2009a) If the pre-fabricated physical voxels self align relative to their neighbors upon assembly, (such as is the case with LEGOsTM or the spherical voxels presented here) then the accuracy of the digital object is purely a function of the precision of the voxels, and not the fabricator that assembled them. This assumes the fabricator is accurate enough to place each voxel within its region of self alignment. Additionally, since the voxels are fabricated separately in highly optimized conditions, they may be composed of any sort of material, and even contain specific functionality.

Additionally, since the voxels are fabricated separately in highly optimized conditions, they may be composed of any sort of material, and even contain specific functionality. For instance, metal, polymer, semiconductor, ceramic, and organic voxels can be created in bulk using processes suitable for each, but in the end, each is poured into the digital fabricator and assembled identically, regardless of material or function as long as it is solid.This opens the door to a desktop process that can print integrated, multi-material 3D objects with high precision. Much like TTL (transistor-transistor logic) standardized the physical parameters (voltage and duration) of digital signals between semiconductor devices, there must be a standard or set of standards defining the interface between voxels in a digital material in order to facilitate widespread adoption. As with other digital technologies, a finite resolution must be chosen based on the needs of the part.

As with other digital technologies, a finite resolution must be chosen based on the needs of the part. To appear smooth to the eye, voxels must be less than approximately 0.5mm. However, larger voxels or smaller may be more suitable, depending on the application, so a spectrum of well-defined sizes may become necessary, such as is the case for screws and other fasteners.Another challenge facing the widespread adoption of digital materials involves the fabrication of the voxels. Technology exists to accurately manufacture voxels at virtually any size scale, such as casting, photolithography, etching, stamping, etc. However, none of these processes are currently cost effective in making millions of voxels. In order for digital materials to be economical, economies of scale must be leveraged to make billions or trillions of voxels in parallel manufacturing processes.

In order for digital materials to be economical, economies of scale must be leveraged to make billions or trillions of voxels in parallel manufacturing processes. Complexity will always have a price, but with appropriate scale and infrastructure, digital materials could conceivably become nearly as cheap as the bulk material itself.Results and DiscussionHere we present the use of a multi-material digital fabricator using 1.5mm spherical voxels to demonstrate the reclamation and re-use of the physical voxels used to manufacture simple objects. Spherical voxels were chosen due to their ready availability in a wide range of materials and their large region of self-alignment upon deposition. This means that the voxels used in this process do not physically interlock, necessitating the use of a binder to hold the voxels together.

This means that the voxels used in this process do not physically interlock, necessitating the use of a binder to hold the voxels together. It is important to note that the binder does not change the underlying principle of the digital material, namely the self-alignment of the voxels, it simply holds them in places during and after assembly.Voxjet Research PlatformThe Voxjet printer used here is shown in Figure 2. This USB-controlled platform is capable of simultaneously printing with three different materials and binding themtogether as they are placed. Additionally, a laser distance sensor enables closed loop feedback to perform error correction on the printed part. The print head scans continuously in a similar manner to an inkjet printer, and can deposit spheres of any given material in a continuous line at approximately 10Hz.

The print head scans continuously in a similar manner to an inkjet printer, and can deposit spheres of any given material in a continuous line at approximately 10Hz. When taking into account the time between rows, motion system, feedback, etc., the printer builds at a net rate of approximately 2-4 voxels per second, depending on the geometry. Thus, a 1000 element structure takes on the order of 5 minutes to build.Figure 2: The VoxJet research platform was used to quickly and autonomously build multi-material structures made of hundreds of 1.5mm spherical voxels. The system also deposits binder to hold the structure together and includes error correction via closed loop deposition feedback.In order to address the challenge of scalability, all the alignment and placing of the spheres is handled mechanically by modular deposition modules.

(Figure 3) To load digital material in the module, spheres of the desired material are simply poured into the reservoir at the top of the voxel deposition module. In the first stage of the printing process, the randomly aligned spheres are ordered into a singe, one-dimensional buffer for printing. Here, the buffer is approximately twenty spheres, which is sufficient to smooth out any random variances in the rate of refilling the buffer under continuous voxel deposition. The buffer is loaded by a continuously rotating paddle wheel at the bottom of the reservoir. Each paddle captures an individual spheres and drops it into the gravity-fed buffer tube. Special care was taken in designing the geometry such that the paddle wheel does not jam with spheres and that spheres are never forced downwards into the buffer. This allows the paddle wheel to rotate continuously, even when the buffer is full, such as in between row depositions.

This allows the paddle wheel to rotate continuously, even when the buffer is full, such as in between row depositions. This eliminates the need for feedback in thisalignment process. The deposition modules were designed modularly such that any number may be stacked side-by side, and a single micro gearmotor attached to one extreme provides the rotation for the alignment wheel of all attached modules.Figure 3: A cutaway diagram of the material deposition modules. Loose spheres are poured into the reservoir, which are aligned into a one-dimensional buffer by the continuously rotating alignment wheel. A solenoid actuated escapement mechanism deposits precisely one sphere with each cycle, and can reliable deposit up to ten spheres per second.The actual deposition of individual voxels is controlled by a variation of an escapement mechanism (Figure 3) actuated by a solenoid. With each cycle of the solenoid, the buffer is advanced under the force of gravity by exactly one voxel.

With each cycle of the solenoid, the buffer is advanced under the force of gravity by exactly one voxel. This works because when the solenoid activates, one tab enters the buffer stream right above the bottom sphere, holding the rest of spheres from advancing while (with the same motion) the tab holding the bottom sphere from falling is removed. When the solenoid is deactivated, it returns to the original position, allowing the buffer to advance by one. Thus, exactly one voxel falls out the bottom of the tube, at a very predictable delay from the actuation of the solenoid.The deposition head also glues and verifies the placement of spheres (Figure 4). The application of the binder is accomplished using the standard method of a pressurized reservoir of binder controlled by a solenoid valve. For these experiments we used a polyvinyl acetate binder in aqueous solution.

For these experiments we used a polyvinyl acetate binder in aqueous solution. Advantages of this system include a relatively quick drying time, mechanical robustness, and the ability to reverse the bonding upon submersion in water. Regular white multi-purpose glue was furtherwatered down at a ratio of 1:1 (glue to water) to yield a solution with low enough viscosity to jet. This was jetted at approximately 10psi (using a regulated diaphragm pump) through a nozzle 0.020" in diameter and solenoid actuation duration of about 17 ms.Closed-loop feedback of the voxel deposition process is accomplished using a Baumer OADM-12 laser distance sensor, which enables error correction on the physical structure. With a theoretical measurement precision of two microns, it is trivial to determine whether a 1.5mm sphere is present or not at any giving location within a layer.

With a theoretical measurement precision of two microns, it is trivial to determine whether a 1.5mm sphere is present or not at any giving location within a layer. The laser sensor works robustly for all materials tested, including those with diffuse surfaces and highly reflective surfaces. Even clear acrylic spheres are sensed correctly, making this a robust, material independent solution for feedback. However, the sensor does not differentiate between materials, so errors involving misplacement of material could pass undetected without the addition of optical feedback.Figure 4: The deposition head consists of a binder depositing system, three separate material deposition modules, and a laser distance sensor. As the print head scans from right to left, binder is deposited appropriately, followed by spheres of any available material.

As the print head scans from right to left, binder is deposited appropriately, followed by spheres of any available material. The laser sensor verifies the placement of each sphere.Recycling multi-material digital materialsWith this robust, repeatable digital material printing framework in place, we demonstrated fully recyclable digital materials. First, two dissimilar materials were chosen to print with, in this case steel and delrin. It may be noted that this combination of materials could not be printed together using any other process. A two-material dome shape was designed for printing, containing approximately 400 voxels. (Figure 5a & b)In the first phase, the center region of the geometry was selected to be delrin, while the periphery region was steel voxels. The structure was printed in approximately three minutes, and allowed to dry for 4 hours at room temperature, to allow the interior ofthe structure to dry.

The structure was printed in approximately three minutes, and allowed to dry for 4 hours at room temperature, to allow the interior ofthe structure to dry. After this, the finished structure was removed from the base-plate. Next the voxels were reclaimed by dissolving the binder. (Figure 5c & d) The entire structure was placed in water and allowed to soak for several minutes. With the help of mechanical agitation, the spheres were separated from each other in solution, then dried. This process was repeated with clean water a second time to ensure minimal glue residue left on the voxels.In the case of steel and delrin spheres, separating them from a common pile is trivial with the help of a low-intensity magnet (Figure 5e). High intensity magnets were avoided so that the steel spheres did not become significantly magnetized, which could interfere in the subsequent alignment and deposition process.

High intensity magnets were avoided so that the steel spheres did not become significantly magnetized, which could interfere in the subsequent alignment and deposition process. The remaining spheres in the reservoir for each material were emptied to ensure that the recycled spheres were in fact the ones being printed, then each constituent material was poured right back into the material reservoir (Figure 5f).The inverse shape was then printed, with steel spheres now in the center region of the geometry and delrin spheres on the peripheral region. This object was then physically printed in a similar amount of time, using the same voxels as the first object (Figure 5g & h). This demonstrates completed recycling of both materials, discounting the trace amounts of binder which was lost in the process.(a)(b)(c)(d)(e)(f)(g)(h)Figure 5: The process of recycling printed material: The initial object is printed consisting of both steel and delrin voxels (a and b).

The finished object is removed from the base plate (c) and dropped into water to dissolve the bonds holding the voxels together (d). After mechanical agitation and drying, the spheres are sorted using a magnet to retrieve the steel spheres (e). The reclaimed raw materials are directly poured back into the deposition modules (f) and another object is printed (g) with the materials reversed (h).ConclusionsHere, we have demonstrated for the first time the complete recycling of the materials used in a multi-material, additive manufacturing process. The process presented here scales effortlessly to arbitrarily large numbers of voxels in an object, arbitrarily small voxels, and an arbitrarily high number of materials, arranged in any spatial combination, including dense interleaving.This paradigm is enabled by the concept of digital material, or matter composed of discrete, self-aligned fundamental building units.

The spheres used here lack geometric interlocking, so binder was used to hold the structure together, which then dominates the properties of the material, especially in tension. Although this binder is not reclaimed here (It certainly could, with some additional effort), the volume of binder is very small compared to the volume of the material in the voxels. By moving to interlocking voxels in the future (Hiller and Lipson, 2009b), the need for binder is eliminated, and the physical properties of the structure are purely a function of the component voxels.Here we demonstrate the recycling of two dissimilar materials which can easily be sorted using a magnet once the structure is decomposed. The challenges in sorting the voxels for re-use becomes more difficult when three or more materials are used, especially when those material have similar material properties. However, we do not expect this to be a significant obstacle to moving forward.

However, we do not expect this to be a significant obstacle to moving forward. Passive sorting based on density will be particularly effective, using vibrations, graded bins, etc., although the process becomes more difficult if two materials with very similar density are used. Efficient active sorting systems also exist which are capable of sorting objects based on optical properties, and one could certainly devise sorting systems based on other material properties such as electrical conductivity, permittivity, dielectric strength, thermal effects, permeability, acoustic properties, etc.Over the course of the last several decades, many technologies have transitioned from analog to digital and never looked back, enabling greater functionality and increased adoption in all cases. Digital materials hold the promise of a precise, multi- material desktop additive manufacturing process with complete recyclability and re- usability of the component voxels.

Digital materials hold the promise of a precise, multi- material desktop additive manufacturing process with complete recyclability and re- usability of the component voxels. Challenges include the need for standardized interfaces between voxels and the infrastructure to produce many voxels at a reasonable cost with interesting functionality. However, when these challenges are addressed, recycling can take place at an unprecedented level. For example, an unneeded camera could be decomposed and turned into a glucose sensor in a desktop process. Waste will be minimized, and a global economy of interoperable voxels could fundamentally change the world of manufacturing.This work was supported in part by a National Science Foundation Graduate Research Fellowship and DARPA Programmable Matter Program, grant W911NF-08-1-0140.References BOURELL, D. L., LEU, M. C. & ROSEN, D. W.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 2(cid:26)th Annual International(cid:3) Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference(cid:3) Reviewed PaperFULLY THREE-DIMENSIONAL TOOLPATH GENERATION FOR POINT-BASED ADDITIVE MANUFACTURING SYSTEMS Maxwell K. Micali*; David Dornfeld*Department of Mechanical Engineering; University of California, Berkeley; Berkeley, CA 94720AbstractWhile additive manufacturing and 3D printing achieved notoriety for their abilities to manufacture complex three-dimensional parts, the state of the art is not truly three-dimensional. Rather, the process plans for the majority of these machines rely on a stack of discretized, two- dimensional layers, which results in parts with stair-stepped surfaces, as opposed to being smooth and freeform.

This work proposes a change to the 2.5D paradigm by using a configuration space approach to enable toolpath planning in a full three-dimensional space, allowing movements beyond planar slices. Algorithms are also presented to compute toolpaths on non-planar regions. Since the toolpaths take part and machine geometries into account, they are guaranteed to be collision-free. These techniques are relevant to many additive manufacturing technologies, including fused deposition modeling (FDM), directed energy deposition (DED), material jetting, and nozzle-based variants. The result of implementing non- planar toolpaths is smoother, more continuous part surfaces.Keywords: CAM, additive, process planning, accessibility, configuration spaceIntroductionAdditive Manufacturing (AM), otherwise known as solid freeform fabrication or 3D printing, has seen a large amount of growth since its inception a few decades ago, both in terms of research activity and industrial adoption.

According to the Wolhers Report, the AM industry was already valued at $3 billion in 2012, and is projected to grow to over $6.5 billion by 2019 [1].There are a variety of AM processes in use, with many more names than there are processes.

In an attempt to standardize the nomenclature of AM processes, the ASTM International Committee F42 on Additive Manufacturing Technologies identified the seven major types of AM processes in its “Standard Terminology for Additive Manufacturing Technologies" in 2009, which are defined as follows [2]:Binder jetting – an additive manufacturing process in which a liquid bonding agent is selectively deposited to join powder materialsDirected energy deposition – an additive manufacturing process in which focused thermal energy is used to fuse materials by melting as they are being depositedMaterial extrusion – an additive manufacturing process in which material is selectively dispensed through a nozzle or orificeMaterial jetting – an additive manufacturing process in which droplets of build material are selectively deposited36(cid:1) Powder bed fusion – an additive manufacturing process in which thermal energy selectively fuses regions of a powder bed(cid:1) Sheet lamination – an additive manufacturing process in which sheets of material are bonded to form an object(cid:1) Vat photopolymerization – an additive manufacturing process photopolymer in a vat is selectively cured by light-activated polymerization(cid:1) Vat photopolymerization – an additive manufacturing process in whichA selection of alternative names to these processes is shown in [2] in order to provide a mapping of some popular process names onto the F42 nomenclature.

The term “rapid prototyping” has historically been one of the most popular terms for AM in general when prototyping was the primary use of the technology. While prototyping still remains an important application, AM technology has progressed to the point where direct part production is possible and increasingly common in a wide selection of build materials.Given the fact that all of these seven core processes merit separate categorization, there are fundamental differences with each process. These differences tend to manifest in the underlying physical mechanisms driving the process, the class of materials available for the process, and the form of the materials prior to the process. However, despite these important distinctions, the digital pipeline from computer-aided design (CAD) to print is identical for all seven AM process technologies.

However, despite these important distinctions, the digital pipeline from computer-aided design (CAD) to print is identical for all seven AM process technologies. The generalized pipeline begins with a digital CAD model, then a slicing algorithm is used to slice the CAD model into a finite number of planar slices along some axis. Information about the slices is sent to a more specialized algorithm which plans how a specific printer will print each slice, then the printer sequentially prints all of the slices into a cohesive structure. It is important to note that the resulting part represents an approximation to the original CAD model as opposed to the exact geometry, since the part is made with a finite number of slices which are in effect sampling the solid model.

More on the topic of planar slicing and its associated issues will be discussed later in this paper.With such variety in AM process technology, it is worth challenging the trope that this generic digital pipeline is optimal for all processes. The intention of this work is to investigate the merits of process-specific pipeline specialization, in hopes that specialization will enable the unique capabilities and benefits of each process to be leveraged in ways which could not be accomplished with a generic solution. More specifically, this work proposes an alteration to the digital pipeline for directed energy deposition, material extrusion, and material jetting additive manufacturing systems, which will henceforth be referred to as “point-based” additive manufacturing systems.

These are point-based technologies because they all have pointed nozzles which eject material; they deposit material selectively and exclusively at the nozzle aperture; and they allow for voids within the total build volume.Conventional ApproachAs mentioned previously, the digital pipeline preceding additive manufacturing processes is typically identical, regardless of the type of manufacturing process being used. The pipeline starts with a 3D solid model, which is then discretized into a finite number of flat slices with the same axial orientation, effectively reducing the sliced representation of the geometry to an approximation of the solid model. These slices are then printed sequentially by the machine,37yielding a physical part which is also an approximation of the intended solid model.

These slices are then printed sequentially by the machine,37yielding a physical part which is also an approximation of the intended solid model. If differentiation does occur in the digital pipeline for a certain process technology, it is merely in determining how the particular machine at hand can most optimally produce each slice, typically optimized for throughput or process stability.There is a compelling reason for why the solid model processing pipeline for AM systems has little variation, and as such why computer-aided manufacturing (CAM) packages for AM offer little more than slicing algorithms. This reason is rooted in the fact that additive manufacturing systems are capable of producing highly complex and intricate geometries, and slicing algorithms are more computationally efficient with handling complexity than featurizing the part’s surfaces and interior, which could also be complex internal lattices.

In addition, slicing algorithms require little supervision from the user, while more sophisticated CAM approaches such as those for conventional computer numerical control (CNC) machining require an experienced user to make a series of decisions throughout the CAM stage of the pipeline.Another advantage of flat slices comes into play when considering the motion of the machines during the printing process. Most formats of AM systems have a mechanical element traversing over the top surface of the work in progress. By only processing material in flat slices, it guarantees that the traversing machine element will not collide with the work underneath, since any given slice has a known maximum height. Changes in the machine element’s height are monotonically increasing, and it is always higher than the highest slice at that stage of the process.

Changes in the machine element’s height are monotonically increasing, and it is always higher than the highest slice at that stage of the process. Collision-free movement is important for any CNC process, since any collision has the potential to damage the machine and/or the part.Conveniences of slicing aside, it also comes with its set of drawbacks. The first drawback is that the height of the machine element is not changed during any given slice, which means AM machines are only 2.5D processes. The 2.5D nature of the process results in quantifiable error when compared to the intended solid model. Error can be minimized with thinner slices, however there is a finite limit to how thin a slice can be. This limit is based on machine resolution or unrealistic manufacturing times, and since this limit always exists, slicing error will always exist whenever there are curved surfaces along the slicing axis.

The error appears as the familiar stair- stepped pattern on the surface of additively manufactured parts, which is often an undesirable aesthetic quality, and results from either under filling or overfilling material with respect to the design surface.Further complicating the use of flat slicing algorithms is the entry of new, 5-axis point- based AM machines in the market. In addition to movement in the x-, y-, and z-axes, they allow for the rotation of the build platform relative to the z-axis, as well as the tilting of the nozzle axis relative to the build platform. In a system possessing these additional degrees of freedom, the question about how to slice the solid model in order to take advantage of the system’s extra capabilities becomes nontrivial, and if that question is not asked, the additional degrees of freedom add little value to the process.

This question can also be interpreted as an opportunity to move beyond flat slices altogether and implement something more sophisticated. Figure 1 shows a summary of the current process and the need to implement new solutions. Figure 2 uses a diagram to demonstrate two types of error resulting from slicing algorithms representing a solid model with a curved surface.38Figure 1: This figure highlights the crossroads faced by AM, where the current, flat slicing processes have fundamental limitations, especially when considering the requirements for systems with more than three axes of movement.Figure 2: A diagram representing two types of error which can result from slicing a solid model with curved surfaces. The regions shaded in red show where the slice is locally undersized compared to the design surface, and the regions shaded in purple show where the slice is oversized compared to the design surface.

The regions shaded in red show where the slice is locally undersized compared to the design surface, and the regions shaded in purple show where the slice is oversized compared to the design surface. The red errors result in ‘gouges’, while the purple errors result in depositing material in regions beyond the design surface.39There have been widespread efforts to optimize particular qualities of additively manufactured parts and the build processes through algorithmically selecting a part’s most ideal orientation on the build plate prior to slicing, as depicted in Figure 3.

The following is an incomplete list of traits which are functions of a part’s orientation:(cid:1) •(cid:1) •(cid:1) •(cid:1) •(cid:1)the total time required to build the part the amount of support material required and its contact area with the part the contact area between the build plate and the part the mechanical properties of the part resulting from the layering direction the total amount of error resulting from slicingFigure 3: Part orientation on the build plate can impact qualities of both the part and the process. Shown here is a simplified 2D projection of a part (red) being built in two different orientations. Since increases in height along the Z axis take longer to print, the orientation on the left takes less time to print even though the part geometry is identical.

Since increases in height along the Z axis take longer to print, the orientation on the left takes less time to print even though the part geometry is identical. In addition, the part on the right requires more support material (blue) in order to build the overhanging structure.It is common currently for slicing programs to include a function to calculate the best part orientation to optimize one or several of the first three bullets listed above. While there have been some research efforts to select the orientation in order to minimize total volumetric error of the slicing direction, these approaches are still affected by issues inherent to parts composed of flat slices, such as highly anisotropic mechanical properties [3].Discretizing a solid model into nonplanar slices rather than flat slices could improve the total volumetric error, as well as allow more control over the anisotropy of the resulting structure.

In order to test these hypotheses, there must first be a way to program an AM machine to follow nonplanar slices without collision. The research detailed in this work addresses the question of how to move toward fully three-dimensional toolpaths for point-based additive manufacturing systems and away from the current 2.5D flat slicing approach to process planning.Relevant LiteratureThis research began by taking a step back and viewing additive manufacturing processes under a more general lens, as opposed to associating them with slicers by default. It became40obvious that point-based AM systems have many features in common with CNC milling machines. The most salient similarities are how the machine is given instructions, how the machine moves, and how well-defined and rigid the tool is inside the machine. Following, strategies used for toolpath generation in the field of CNC machining can be relevant literature for fully three-dimensional additive manufacturing toolpaths.

Following, strategies used for toolpath generation in the field of CNC machining can be relevant literature for fully three-dimensional additive manufacturing toolpaths. For the remainder of this paper, it is assumed that “additive manufacturing" is referring to point-based additive manufacturing systems unless otherwise specified.Another similarity between AM and CNC milling can be seen when comparing the surfaces of a printed part and a part which has been through roughing cuts on the mill. If the design geometry is curved, both part surfaces will have a stair-stepped pattern. In this case the milling machine is also operating on geometry which has been discretized into slices – however, the surface does not remain this way in a completed milled part. Roughing is just a precursor to the finishing passes for machining complex, freeform surfaces.

Roughing is just a precursor to the finishing passes for machining complex, freeform surfaces. While the roughing operations may operate in 2.5D, the finishing passes trace the contours of smooth surfaces, requiring fully three-dimensional toolpaths. This process of a milling cutter tracing intricate, smooth surfaces is called Sculptured Surface Machining (SSM), and it is a technique commonly used in machining die-cavities. Methods used in generating finishing toolpaths for SSM are important to consider in generating toolpaths for nonplanar AM.SSM is a point milling process where the milling cutter moves through a series of cutter contact (CC) points, which are points where the cutter is in contact with the design surface. Another way to parameterize the positions of the cutter is to use cutter location (CL) points. CL- points are unique for each position of the cutter, and are defined based on some reference point on the tool, such as the spherical center of a ball nose end mill.

CL- points are unique for each position of the cutter, and are defined based on some reference point on the tool, such as the spherical center of a ball nose end mill. When the tool is slid over the design surface of the part, the CL-points form a CL-surface. A diagram of the CL-surface is shown in Figure 4. Since the tool cannot be slid over an unmachined stock surface to form the CL-surface for generating a design surface, other methods must be used to calculate the CL- points a priori.The CL-points can be calculated locally by using a CC-point approach or a direct positioning approach. The CC-point approach first plans the toolpaths on the design surface, then computes the CL-points from that set of CC-points. Sata et al. [4] used an isoparametric method to plan the CC-points, but difficulties arose with the non-uniform relationship between the parametric coordinates and the physical coordinates [5]. Huang et al.

[4] used an isoparametric method to plan the CC-points, but difficulties arose with the non-uniform relationship between the parametric coordinates and the physical coordinates [5]. Huang et al. [6] proposed an algorithm based on nonconstant parameter tool contact curves, and Suresh and Yang [7] proposed iso- curvature methods along characteristic lines. A direct positioning approach can also be used, which plans 2D toolpaths first on the horizontal guideplane, then lowers individual tool positions until they touch the design surface [8].41Figure 4: Diagram showing the construction of a CL-point and a CL-surfaceLocal toolpath generation methods are not ideal because they are prone to gouging. Two locally safe CL-points could be linked to form a CL-line which is not a gouge-free path over the design surface, and this is difficult to anticipate using local methods. In the context of high speed machining, any gouge or collision could be disastrous.

In the context of high speed machining, any gouge or collision could be disastrous. As a result, the needs for high speed machining drove Choi et al. [8] to develop a configuration space (C-space) approach to toolpath generation, which is a global approach and therefore less prone to gouging and collisions than local methods.C-spaces are a concept used in robotic spatial planning problems, especially for automatic manipulator movements [9]. Each element of the C-space corresponds to a valid system configuration, and these configurations can be described by whether or not they are safe. Safety is defined as the set for object A and obstacle B where A ∩ B = Ø. Applying C-spaces to cutting tools and the problem of toolpath generation yields some convenient properties, since it guarantees gouge-free and collision-free paths. In order to accomplish this, the first step in the toolpath generation process is to transform the design surface into C-space elements.

In order to accomplish this, the first step in the toolpath generation process is to transform the design surface into C-space elements. Once this is done, the toolpaths are generated within the C-space elements rather than on the design surface, which automatically incorporates information from the global part geometry in the path planning process.A convenient C-space transformation to use is the inverse tool offset (ITO), shown in Figure 5. The ITO is constructed with the Minkowski sum concept. This works out in practice by inverting the tool, locating the CL-point of the inverted tool on the design surface, then sweeping out an envelope of the inverted tool’s geometry while moving along the design surface. The envelope is the ITO-surface, and it also defines a safe C-space for the tool. During operation, the ITO-surface is the CL-surface required in order to machine the design surface.

The envelope is the ITO-surface, and it also defines a safe C-space for the tool. During operation, the ITO-surface is the CL-surface required in order to machine the design surface. Thus, toolpaths can be generated within the ITO-surface to have effective, yet gouge-free and collision-free operation.It is worth noting some key differences about AM which must be taken into account if a similar ITO method is to be used in a C-space approach to toolpath generation. With CNC42machining, it is assumed that if a cutter contacts a surface, the cutter will machine the surface, provided the CC-point(s) are on a cutting surface of the tool for a given CL-point. In addition, contact with a cutting surface on the tool is sufficient to consider machining to have occurred. In AM on the other hand, which will be assumed to have nozzles which can be modeled with a right cone, only a very specific point of the nozzle can deposit material – the tip.

In AM on the other hand, which will be assumed to have nozzles which can be modeled with a right cone, only a very specific point of the nozzle can deposit material – the tip. Thus if the tip is not in contact with the design surface but the side of the nozzle is in contact as a result of the ITO, this is not contributing anything to the deposition process. Rather, it is a sign that the tool has approached its limitation in that region of the part. Also it is not enough to simply program the position of an AM tool, since the material deposition process must be programmed as well.Figure 5: Inverse tool offset (ITO) surfaceXY-Parallel Toolpath Generation for Three-Axis, Point-Based Additive Manufacturing SystemsChoi et al. in [8] proposed an XY-parallel one-way finishing unit machining operation (UMO) algorithm as a proof of concept for taking a C-space approach to SSM.

in [8] proposed an XY-parallel one-way finishing unit machining operation (UMO) algorithm as a proof of concept for taking a C-space approach to SSM. A logical extension of that algorithm with appropriate enhancements and modifications can serve as a proof of concept for applying these approaches to AM toolpath generation. That proof of concept is the subject of this work. Figure 6 shows a schematic of the overall toolpath generation algorithm for XY-parallel toolpaths for three-axis AM systems. It begins with the user specifying a minimum number of parameters: the design surface geometry, the geometry of the nozzle, the tolerance of the process or the minimum resolution of the machine, and the maximum step- length allowed along a path. The algorithm proceeds first by generating an ITO-surface from the design surface and the nozzle geometry.

The algorithm proceeds first by generating an ITO-surface from the design surface and the nozzle geometry. Since the ability of conical nozzles to access narrow regions decreases with the aspect ratio of the nozzles’ height to its base diameter, the printability of the part is a function of whether a given nozzle can access all of a given design surface. Unless the entirety of the design surface is accessible, the part cannot be printed as intended. To address this issue more concretely, a printability check is incorporated in the algorithm after the ITO-surface is generated. If the part is deemed unprintable, the algorithm will inform the user that the aspect ratio of the nozzle is too low for the given design surface. Otherwise, the process will proceed to the next step of defining XY-parallel paths within the C-space, followed by calculating control commands for the material deposition function of the machine.

Otherwise, the process will proceed to the next step of defining XY-parallel paths within the C-space, followed by calculating control commands for the material deposition function of the machine. All of this information is finally synthesized by a script that converts it into a standardized G-code format43which AM systems can understand, provided the proper machine-specific heading is used at the top of the file. Since few AM machines support circular interpolation commands (G02 & G03), the types of output commands for this toolpath generation algorithm are limited to linear interpolation (G01) and rapid traverse (G00).Figure 6: Schematic of XY-parallel toolpath generation algorithm for three-axis AM systemsThe design surface used to test these algorithms is shown in Figure 7 [10]. The format of this surface is a Z-map, which is a matrix of height values in an equally spaced grid corresponding to the X and Y indices of the matrix.

The format of this surface is a Z-map, which is a matrix of height values in an equally spaced grid corresponding to the X and Y indices of the matrix. The doubly curved nature of this surface makes it particularly apparent whether the toolpaths can trace the surface curves.44Figure 7: Design surface used to test algorithmsGenerating the ITO for a nozzle modeled as a right cone proceeds similarly to how it does for other tools seen in CNC machining. It starts by inverting the cone, then placing the nozzle location (NL) point along the design surface, sweeping out an envelope of the nozzle body during the process. A diagram of this can be seen in Figure 8 below. Terminology is changed to NL-point from CL-point to be more consistent with the process, since this nozzle does not cut any material. The red cones are inverted nozzles, and the purple shaded regions depict the portion of the ITO-envelope which exists above the design surface.

The red cones are inverted nozzles, and the purple shaded regions depict the portion of the ITO-envelope which exists above the design surface. The nozzle in this figure has too low of an aspect ratio to access this entire surface, as shown by the blue triangles in the figure which are at the edge of the allowable envelope but do not have their tips in contact with the design surface. In order to reach the design surface with the nozzle tip in these regions, the nozzle would need to intersect with the surface, and this is not an allowable configuration in the safe C-space. Since the sides of the cone do not deposit material, this nozzle would not be able to print this design surface.45Figure 8: (left) Diagram of nozzle and its nozzle location (NL) point. (right) Diagram of the ITO- surface generation process with a right cone tool geometry on an arbitrary design surface. This cone’s aspect ratio is too low to fully access this design surface.

(right) Diagram of the ITO- surface generation process with a right cone tool geometry on an arbitrary design surface. This cone’s aspect ratio is too low to fully access this design surface. The purple regions indicate the zones inaccessible by the nozzle tip.It is possible to calculate whether a given nozzle geometry would be capable of printing the entirety of a given design surface, and one approach to calculating this is incorporated within the present algorithm. Since the NL-point corresponds to the tip of the nozzle and this point must be able to access any point of a design surface for it to be fully printable, any separation between the ITO-surface and the design surface which is greater than the manufacturing tolerance at any given point can render that nozzle/surface combination “unprintable". The ITO in this application has the most value for being a convenient method for algorithmically determining printability.

The ITO in this application has the most value for being a convenient method for algorithmically determining printability. See Figure 9 for a visual depiction of the NL-surface and design surface delamination due to a nozzle aspect ratio error. A field showing the local magnitude of separation between the two surfaces is plotted beneath the surfaces. From the magnitude field and the NL-surface, it can be deduced that there is a maximum allowable slope for the movement of the nozzle within the safe C-space, which is consistent with what would be expected for a vertical cone traversing a surface. The solution to a nozzle aspect ratio error like the one shown is to use a nozzle with a higher aspect ratio in order to enable the manufacturing of steeper surfaces.

The solution to a nozzle aspect ratio error like the one shown is to use a nozzle with a higher aspect ratio in order to enable the manufacturing of steeper surfaces. Even with an infinite aspect ratio nozzle, three-axis AM systems will not be capable of producing a vertical feature within a single, curved layer.46Figure 9: Delamination of the NL-surface and design surface due to a nozzle with insufficient aspect ratio for the given design surface. The NL-surface and design surface are plotted on top of each other, and a field showing the local magnitude of their difference is plotted below them.Once the ITO has been generated and the printability is verified, it is time to generate XY-parallel toolpaths. Since the toolpaths are calculated within the safe C-space, it is guaranteed to be collision-free. The algorithm implemented for this stage is adapted from the work done by Choi et al. for the XY-parallel paths in [8].

The algorithm implemented for this stage is adapted from the work done by Choi et al. for the XY-parallel paths in [8]. Pseudocode for this algorithm is below, where ω is the path width, λ is the maximum step-length, α is an inclination angle in either the feed or pick direction, and tol is the manufacturing tolerance.

In practice, ω is likely to be close to the aperture diameter of the nozzle, and it makes sense to set tol equivalent to the resolution of the AM machine.Start at edge, Δy = ωOutput NL-point (X,Y,Z) Fit circle in feed-forward direction λ = min(λmax, 2*sqrt(2*tol*Rf)) Calculate inclination angles αf,p, in feed & pick directions x = x + λ*cos(αf) Δy = min(Δy, ω*cos(αp)) Continue until opposite edgeNew pass: y = y + Δy, x = xo, Δy = ω Continue until done Linearly interpolate to form NL-lines47With a set of NL-lines and NL-points, it is possible to calculate commands to send with the G-code output to the part of the control system which manages the material deposition in the AM machine, such as a powder feeder or extruder motor. The following analysis will proceed as if it is an extruder motor in a material extrusion system, as shown in Figure 10, but this analysis could be reformulated to cater to a powder feeding system in a directed energy deposition system.

Since the aperture at the nozzle tip is smaller than the diameter of the incoming filament, the change in linear position due to the extruder motor on the filament is amplified upon exit of the nozzle. All steps along NL-lines are linear interpolations because many AM system controllers cannot interpret higher order commands, so it is appropriate to calculate the path length between any two successive NL-points by calculating their three-dimensional Euclidean distance. Once the nozzle’s path length is calculated, the required motion of the extruder motor can be calculated to satisfy that extruded path length, and the current position of the motor is updated. These calculations are shown below in Equations 1, 2, and 3 for a span between two arbitrary points i and j. E refers to the absolute position of the extruder motor, and d refers to the diameter of the given end of the nozzle. Please refer to Figure 10 for explanations of the variables.

The coarse step-length is ten times the maximum step-length of the fine step-length. These outputs step along the x-axis, however stepping could be performed along any direction. Near the uppermost portion of the surface, all of the parallel paths start identically. As they travel down the slope and the surface becomes more doubly curved, the paths begin differentiating and adapting to the local curvature. This is easiest to see in the coarse figure. Both surfaces look like the design surface in Figure 7, and the two step-length trials pictured produce surfaces which look similar. There must be some ideal maximum step-length for each design surface and nozzle combination, after which increasing resolution via step-length does not improve the quality of the manufactured part.

There must be some ideal maximum step-length for each design surface and nozzle combination, after which increasing resolution via step-length does not improve the quality of the manufactured part. This leads to the question of how to automatically determine the maximum tolerable maximum step-length, however that topic is saved for future work.Figure 11: Two rotated renderings of sample output for the test surface with steps along the x- axis. Grid spacing of the Z-map is 0.0141 arbitrary units, and the maximum step-length is 0.1 arbitrary units. Blue dots show NL-points, and red dashed lines show interpolated NL-lines.49Figure 12: Two rotated renderings of sample output for the test surface with steps along the x- axis. Grid spacing of the Z-map is 0.0141 arbitrary units, and the maximum step-length is 0.01 arbitrary units. Blue dots show NL-points, and red dashed lines show interpolated NL-lines.To see how the extruder motor behaves over time, please refer to Figure 13.

Blue dots show NL-points, and red dashed lines show interpolated NL-lines.To see how the extruder motor behaves over time, please refer to Figure 13. The position of the extruder is monotonic increasing, which indicates that the algorithm is performing as expected. In addition, the plot of position vs. NL-point shows variability in the extruder position update, as expected since the path length with each step between NL-points is variable. There are occasionally two NL-points which have identical extruder motor positions, and this is when the nozzle traverses from the end of one parallel path to the beginning of the next one, so there should be no extrusion during this traversal.Figure 13: (left) Rendering of extruder motor absolute position shown in color at each NL-point for a Z-map with grid spacing of 0.0141 arbitrary units and maximum step-length of 0.1 arbitrary units. The position value increases from dark blue to yellow.

The position value increases from dark blue to yellow. (right) A plot of the extruder motor position for the first 100 NL-points.Conclusions and Future ResearchThe results of this work are algorithms to automatically generate of XY-parallel toolpaths for point-based additive manufacturing systems. This is one of hopefully many examples of the digital pipeline being tailored to specific additive manufacturing processes in order to leverage the unique strengths they each afford. Beyond that, these results are important because they show that a third dimension can be activated in 3D printing, and they also show that the field of additive manufacturing can draw from the toolpath generation work that has been done in the field of CNC machining with careful and thoughtful modification.The C-space approach seems to be most fruitful as a method for determining printability for a given design surface, nozzle geometry, and manufacturing tolerance.

Calculating the ITO- surface is a convenient way to determine printability, and it is important to know whether a part is printable before attempting.50By using fully three-dimensional toolpaths, the printer nozzle can trace the complex contours of a curved surface as opposed to discretizing them and approximating them with a stair-stepped appearance. This will certainly reduce the total amount of volumetric error in printed parts when compared to the intended solid model. While very steep curves are difficult even with high aspect ratio nozzles, there will be a wide set of parts that can still benefit from using this approach. This method allows a printer to effectively increase the resolution of its products without modifying the components of the printer, since this set of algorithms outputs G- code.

This method allows a printer to effectively increase the resolution of its products without modifying the components of the printer, since this set of algorithms outputs G- code. Another benefit of converting instructions to G-code in the algorithm is that, since many AM system controllers can interpret G-code, the methods contained in this paper can benefit all of those printers which are currently in operation. While there would be more opportunities for interesting extensions to this work if machine controllers could interpret G02 and G03 circular interpolation commands, this is a hurdle which is difficult to overcome for existing machines.Some future research projects which are in progress involve validating these algorithms on AM systems, which includes characterizing the surfaces of the final part, as well as the reduction in total volumetric error.

It would be interesting to investigate how nonplanar slicing affects the mechanical properties of a part when compared to flat slicing algorithms.The emphasis of this work is on the geometric path itself, however the machine dynamics and the underlying physical phenomena in additive manufacturing systems are also critical for successful part production. The author is looking forward to incorporating physical modeling into the toolpath generation process in future iterations and extensions of this work.AcknowledgementsThis material is based upon work supported by the National Science Foundation Graduate Research Fellowship Program under Grant No. 1106400. Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the views of the National Science Foundation.[1] “Wohlers Report 2012. Additive Manufacturing and 3D Printing State of the Industry,”Wohlers Associates, Ft.

NCTIONAL GRADIENT MATERIAL OF Ti-6Al-4V AND γ-TiAl FABRICATED BY ELECTRON BEAM SELECTIVE MELTINGWenjun Ge*†, Feng Lin*†, Chao Guo*†Key Laboratory for Advanced Materials Processing Technology, Ministry of Education of China, Tsinghua University, Beijing 100084, P.R. China †Center for Bio-manufacturing and Rapid Forming Technology, Department of Mechanical Engineering, Tsinghua University, Beijing 100084, P.R. ChinaAbstractin fabricating functionally graded materials. Electron Beam Selective Manufacturing (EBSM) is one widely used AM technology capable of fabricating a variety of materials especially titanium alloys. Previous studies on EBSM process were focused on the manufacturing of one single material. In this study, a novel EBSM process capable of building gradient structures with dual metal materials was developed. Ti6Al4V powders and Ti47Al2Cr2Nb powders were used to fabricate Ti3Al/TiAl and Ti6Al4V/Ti3Al dual metal structures.

Ti6Al4V powders and Ti47Al2Cr2Nb powders were used to fabricate Ti3Al/TiAl and Ti6Al4V/Ti3Al dual metal structures. The chemical compositions, microstructure and micro-hardness of the dual material samples were investigated employing Optical Microscope (OM), Scanning Electronic Microscope (SEM), Electron Probe Micro-Analyzer (EPMA). Results showed that the thickness of the transition zone was about 300μm. The transition zone was free of cracks, and the chemical compositions exhibited a staircase-like change. The microstructure and chemical compositions in different regions were studied. Micro- hardness was affected by the microstructure.

The microstructure and chemical compositions in different regions were studied. Micro- hardness was affected by the microstructure. The microstructures turned out to be full lamellar at the TiAl region and basket-weave structure at the Ti3Al and Ti6Al4V region.Additive Manufacturing (AM)technologies are very promisingIntroductionTitanium (Ti) alloys are widely used in the aerospace and aviation industries because of their excellent properties including low density, high specific strength, excellent corrosion resistance in various kinds of environment and good high-temperature performances. Ti6Al4V have been widely studied but the working temperature is relatively low. Both the TiAl and Ti3Al- based alloys are attractive materials for high temperature structural applications in aerospace industry due to their high specific strength and good oxidation resistance at elevated temperature [1~4].

However, they suffer from poor ductility at the room temperature and low fracture toughness, which prevent their commercial exploitation [4].Functional Gradient Material (FGM) shows great potentials for the applications in aerospace, dynamic high-pressure technology and medical implants [5]. Specifically, in the aerospace industry, compressors and turbine disks, which are no doubt the most significant parts, require different material properties at different regions. For instance, the rim regions, which suffer from high temperature and relatively small centrifugal force, mainly require good oxidation resistance. On the other hand, the core regions are faced with challenges of extremelylarge centrifugal force and relatively low temperature, and require high strength rather than good oxidation resistance.

On the other hand, the core regions are faced with challenges of extremelylarge centrifugal force and relatively low temperature, and require high strength rather than good oxidation resistance. Yang [6] reported that the rim of a dual alloy disk can be made of intermetallic such as Ti3Al or TiAl alloy with good high-temperature properties, and the core can be made of titanium alloy such as TC4 or TC11 with good properties at relatively low temperatures.Additive Manufacturing (AM) technologies are quite promising in the fabrications of FGM. Most of the current additively manufactured FGMs were fabricated by laser-based additive manufacturing technologies e.g. Selective Laser Melting (SLM). Xiehang [7] fabricated the Ti6Al4V-CoCrMo thin wall gradient material through laser direct forming and investigated the effect of CoCrMo volume fraction on the cracking behaviors.

Xiehang [7] fabricated the Ti6Al4V-CoCrMo thin wall gradient material through laser direct forming and investigated the effect of CoCrMo volume fraction on the cracking behaviors. The main reasons of the fracture are that the limited solid solution network CoCrMo caused the stress concentration and the mismatch of thermal properties. Xi [8] investigated laser rapid forming of 316L stainless steel/Ni-based alloy/Ti6Al4V gradient materials. The phase change in the gradient segment of Ni-based alloy and Ti6Al4V was studied and they found that the hardness values decreased with the increase of the content of Ti6Al4V. Xu [9] fabricated TC11/γ-TiAl bio-materials using laser deposition process. According to his study, the compositions insignificantly affected the cracks in the transition zone. The microstructure transformation mechanisms of TiAl and TC4 were also investigated. Douglas C.

According to his study, the compositions insignificantly affected the cracks in the transition zone. The microstructure transformation mechanisms of TiAl and TC4 were also investigated. Douglas C. Hofmann [10] fabricated Ti6Al4V-V gradient metal alloys through laser-based radial deposition additive manufacturing.However, one main problem of the laser-based AM technologies is the fracture and the oxidation. Electron Beam Selective Melting (EBSM) has several advantages over laser-based AM technologies in functional gradient titanium alloys. First, preheating the base plate and the powder bed to a high temperature could reduce the thermal stress, especially for the brittle materials like TiAl and Ti3Al. In addition, the vacuum environment could prevent the introduction of impurity elements. In this work, we fabricated FGMs employing a novel EBSM system developed by Tsinghua University. TiAl/Ti3Al and Ti6Al4V/Ti3Al FGMs were designed and fabricated.

In this work, we fabricated FGMs employing a novel EBSM system developed by Tsinghua University. TiAl/Ti3Al and Ti6Al4V/Ti3Al FGMs were designed and fabricated. The Ti3Al was obtained by mixing the two metal powders of TiAl and Ti6Al4V. The two powders can form homogenous chemical compositions in the melting process. The chemical compositions, microstructure and micro-hardness were studied.Material and experiment procedureTwo kinds of titanium alloy FGMs were fabricated through EBSM. The EBSM system used in this study was developed by the Center for Bio-manufacturing and Rapid Forming Technology in the Department of Mechanical Engineering, Tsinghua University. Figure.1 shows the schematics of the supplying system of dual powders [11]. This system contains vibrator, two powder tanks and powder-mixing box.

Figure.1 shows the schematics of the supplying system of dual powders [11]. This system contains vibrator, two powder tanks and powder-mixing box. Once the value of two powders has been set, the vibrators vibrate to make the powder flow and dropped into the powder-mixing box precisely and then the powder-mixing box rotates back and forth several times to make the composition homogenous. Finally, the mixed powders were spread onto the base plate.Atomized Ti6Al4V and Ti47Al2Cr2Nb powders were used to fabricate Ti6Al4V/Ti3Al and TiAl/Ti3Al gradient materials. Ti3Al was obtained through the mixing of Ti6Al4V and Ti47Al2Cr2Nb. Figure.2 schematically shows the two samples.(a) powder mixing(b) powder spreading Figure.1 Powder supplying systemThe baseboard of 316L stainless steel with the dimensions of 90mm× 90mm×10mm, was grounded and washed in ethanol to remove surface oxidation film, before it was placed into the vacuum chamber.

The substrate was firstly preheated to 800℃using a defocused electron beam. Besides, during the fabrication process, each 100μm thick layer of blended powders was also preheated to be lightly sintered employing the beam current ranging from 1mA to 15mA and the scanning velocity of 5m/s. For each layer, the compositions are affected by both the fabrication parameters and the compositions of the following deposited layers. In order to investigate the effects of the fabrication parameters on the transition zone, the samples were fabricated with different beam currents as shown in Table.1 (wt% denotes the mass percentage). For example, every layer of Sample-1 was scanned to be melted for 3 times with the increasing beam currents from 4mA to 8mA, and then to 12mA.

Then a new layer of powders was spread on top of theprevious layer, and the preheating and melting process were repeated until the sample was finished.(a)sample-1 (b)sample-2 Figure.2 Composition model of dual metal materialsThe samples were cut along the building direction using Wire Electrical Discharge Machining and then mounted, ground, polished and etched utilizing Kroll's reagent (2% HF, 5% HNO3 and 93% H2O).Micro-hardness was measured on MH-3microhardness tester with the load of 200g and the dwelling time of 15s.Microstructures of EBSM fabricated components were observed via Optical Microscope (OM) and Scanning Electron Microscope (SEM).Semi- quantitative and quantitative analyses of the chemical compositions were performed using Electron Probe Micro-Analyzer (EPMA).Table.1 Compositions and fabrication parameters of each layer Layerwt% TC4 wt% TiAlBeam CurrentScanning speed 0.2m/s0-100%100%2mA-6mASample-110-2050%50%4mA-8mA-12mA0.5m/s0-30100%0%2mA-4mA-6mA0.5m/sSample-230-4550%50%2mA-4mA-6mA0.5m/s45-5040%60%2mA-4mA-6mA0.5m/sResultsChemical compositionsTo avoid the delamination, the previously formed layer needs to be at least partially re- melted.

The re-melting depth would also significantly affect the composition gradient of the transition zone. Therefore, the compositions of the transition zone could be significantly affected by the fabrication parameters such as electron beam current, scanning speed and layer thickness. Semi-quantitative and quantitative analyses of the as-built samples were performed employing EPMA. Figure.3 shows the top region of the transition zone of Sample-1. As shown in Figure.3, there were no cracks or un-melted powders and the chemical compositions were relatively uniform in the transition zone. There were two zones with different chemical compositions and both the heights were 100um which is the same with the layer thickness. The compositions of thetransition zone varied stepwise, due to the layer-fashioned manufacturing process. EPMA analysis results showed that element change in the transition gradient zone.

The compositions of thetransition zone varied stepwise, due to the layer-fashioned manufacturing process. EPMA analysis results showed that element change in the transition gradient zone. When the first layer of Ti3Al was deposited on top of the Ti6Al4V or TiAl, a portion of Ti6Al4V or TiAl was re- melted and mixed with the melted powders of Ti3Al in the molten pool. According to the distribution of elementary Al and Ti, the re-melting depth of TiAl could be calculated to be about 200μm.Figure.3 Chemical composition of transition region of Sample-1For the Ti6Al4V-Ti3Al sample, there are two transition regions. Figure.4 shows the surface scan results of the top and middle transition regions. The results indicated that the top layer as the final deposition layer had a depth of 180μm. Considering that the powder layer was 100μm thick, it can be concluded that the height of the re-melted layer thickness was 80μm.

Considering that the powder layer was 100μm thick, it can be concluded that the height of the re-melted layer thickness was 80μm. The results of the line scan showed the distribution of elementary Al and Ti in the transition zone asFigure.5. The depth of the middle transition zone was 400μmand the content of Ti and Al changed gradually. In the top transition region, the curve is flatter except the top layer. Compared with the Ti3Al-TiAl gradient zone, composition changed more smoothly in the Ti3Al- Ti6Al4Vtransition zone. No step shapes could be seen on the curve. The reason is that the composition difference between Ti3Al and Ti6Al4V is lower than the difference between TiAl and Ti3Al.

No step shapes could be seen on the curve. The reason is that the composition difference between Ti3Al and Ti6Al4V is lower than the difference between TiAl and Ti3Al. Also, the re-melting depth h is smaller resulting in a smooth change.(a)top transition region (b) middle transition regionFigure.4 Surface scan resultsTiAl010020030040050060070080090001020304050607080Distance(μm) Conc（%）AlConc（%）Distance（μm）01002003004005006007008009000102030405060708090Ti(a)top transition region (b) middle transition regionFigure.5 Liner scanning resultsMicrostructure evolutionsThe properties of gamma TiAl alloys depend strongly on the grain size and the thickness of γ-TiAl and α2 –Ti3Al lamellar. Refined full lamellar microstructures are most attractive, since they offer controllable strengths, good creep and fatigue resistance, good fracture toughness and reasonable room temperature ductility.

As can be seen in Figure.6 (d) ,the microstructure of TiAl part was full γ-TiAl and α2-Ti3Al lamellar. The size of the lamellar cluster was 15~30μm, which is finer than as-cast TiAl alloy full lamellar microstructure(200μm).As in the as-built sample, a refined lamellar colony of 15μm in size was found along the whole TiAl part.In the Ti3Al region, EPMA quantitative analysis results showed that Al content was 26wt.%. According to Ti-Al phase diagram [12], α phase transformed to α2 phase. The phases at Ti3Al region consist of α2, β/B2 forming basket-weave structure. Figure.6 (b) and (c) show the microstructure of transition zone and the top zone. It can be observed that in the transition zone the width and length of α2 are 0.5μm and 5μm respectively, while in the top zone the width and length are 1μm and 10μm respectively.Figure.6 microstructures of sample-1Figure.7shows the OM images and SEM images of Sample-2.

The images show that the microstructures of Ti6Al4V part are of typical EBSM fabrication features. Microstructures of EBSM produced Ti6Al4V alloy consist of large columnar prior β grains( see Figure.7(c)) growing along the building direction and basket-weave(α+β) structure and widmanstetten α platelets(see Figure.7(d)). The width of prior columnar β grains was about 100μm. Fine α platelets are precipitated along the βgrain boundary and inside the grains. It should be noted that the average α platelets thickness was 2μm forming basket-weave and widmanstetten structure.The microstructures of Ti3Al part can be seen in Fig.7 (a) and (b). It can be observed that the α2 phase exhibits long lathes shape on the top region in Fig.7 (a) while it exhibits short-rod like shapes in Fig.7(b) in the middle region.

It can be observed that the α2 phase exhibits long lathes shape on the top region in Fig.7 (a) while it exhibits short-rod like shapes in Fig.7(b) in the middle region. The thickness of α2-laths was 1μm which was finer than Ti6Al4V region.(a) (b)(c) (d) Figure.7 Microstructures of Sample-2Micro-hardnessFigure.8 shows the Vickers micro-hardness of the dual alloy specimens measured from Ti3Al region to TiAl region. Since the deposition layer thickness was 100μm, the interval between measuring points was 50μm. From the result, it can be observed that from the Ti3Al region to TiAl region the fluctuations are small. Its maximum value is up to 450MPa at the Ti3Al region. The first two points represent the top region of Ti3Al and the micro-hardness is higher than the bottom region of Ti3Al. The reason is that the microstructure at the top region is fine martensitic α laths.

The reason is that the microstructure at the top region is fine martensitic α laths. In the transition zone, the hardness shows a normal level distributing in the range of 350MPa-400MPa.Figure.7 Micro-hardness along the building direction of sample-1DiscussionThe previously fabricated layer needs to be melted and that will make a dilution in composition. The penetration depth was significantly affected by the process parameters such as electron beam current, scanning speed and layer thickness. Based on the re-melting depth the composition change can be calculated.The following simplifying assumptions could be made. (1)In the deposition layer, composition is homogenous. (2)The solid diffusion of element is ignored. The composition of forming layers will notchange.(3)The shrinkage of the powder layer after melting is ignored.

From the result, it can be conclude that the result of tested is well matched the calculation. This validated theoretical model is able to provide both accurate explanations and design disciplines for the fabrications of FGMs.)% wt( lA28 26 24 22 20 18 16 14 12 10calculated value test value123456789Layer Fig.8 Calculated and measured distributions of element AlThe TiAl alloy underwent several thermal cycles. Rapid heating and cooling result in fine lamellar structures. The temperature history during the fabrication process significantly affected the formation of the lamella structures. The EBSM process employs high power density electron beam and the cooling rate could be as high as 103 ~ 104K/s. The high cooling rate associated with the electron beam melting process is the reason for the formation of fine microstructures in the EBSM products. Peng [13] employed the cyclic heat treatments to obtain fine lamellar microstructures.

Peng [13] employed the cyclic heat treatments to obtain fine lamellar microstructures. The grain size could be reduced to 20μm after 7 heat treatment cycles. In the rapid heating process, the transformation occursα+γ→α when the temperature reaches Tα. The nucleation of fine lamellar microstructure mainly occurs along the grain boundaries and also at the phase interfaces. With the increase of the heating rate, the nuclear rate increases and the transfer rate is faster. In the EBSM process, the formed layers will undergo several rapid heating and cooling cycles.The microstructure of Ti6Al4V and Ti3Al are the same, consisting of α/α2 and β. The formation of α2-Ti3Al is controlled by the content of element Al. Al is an α phase stability element. According to Ti-Al phase diagram of Al, when the content is higher than 7%, α2- Ti3Alphase will penetrate. Al is α phase stability element, in the first deposition layer of Ti3Al, the Al element is higher and makes α phase grow up difficult.

Al is α phase stability element, in the first deposition layer of Ti3Al, the Al element is higher and makes α phase grow up difficult. Also the βstability element is alsohigh, makes the portion of β phase is higher. That makes the α/α2 phase become the short acicular shape or the lamellar shape, and their distribution is dispersal too. According to investigate of Yang [6] that the transition zone is the most active diffusion zone, the elements Ti, Al, Cr, Nb and V will diffusion from Ti3Al side and TiAl side to the diffusion zone, that makes α/α2 phase accumulation and grow up difficult. The shape of α/α2 phases become short acicular and the distribution is dispersal.ConclusionTitanium dual metal materials were successfully fabricated using EBSM. The effect of fabrication parameters was analyzed in order to understand the composition change and phase transformations. In addition, micro-hardness of transition zone was analyzed.

The effect of fabrication parameters was analyzed in order to understand the composition change and phase transformations. In addition, micro-hardness of transition zone was analyzed. The following conclusions can be derived from this study.(1)The energy input for Sample-1 is higher than Sample-2 which makes the melting depth larger. According to the composition analysis, the re-melting depth of Sample-1 is 200μm, while the re-melting depth in Sample-2 is 80μm.(2)The microstructure of TiAl region was full γ-TiAl and α2-Ti3Al lamellar structure with colony size 15~30μm. In the Ti3Al and Ti6Al4V regions, the microstructures are similar.

nctional Gradient Metallic Prototypes through Shape Deposition ManufacturingJ. Fessler, A. Nickel, G. Link, F. Prinz Stanford UniversityP. Fussell ALCOAAbstractStanford's SDM laser deposition system has been recently improved to enable the deposition of functionally graded metals through the use of powder mixing. While Shape Deposition Manufacturing has always had the capability to produce multimaterial artifacts, powder mixing enables the deposition of single layers in which material properties can be smoothly varied without discrete interfaces between dissimilar materials. It has been shown that certain materials will completely mix during deposition and form alloys which exhibit material properties intermediate to those of the constituent feed powders. To date, oxidation and hardness have been effectively controlled through appropriate mixing of powders.

To date, oxidation and hardness have been effectively controlled through appropriate mixing of powders. Functional gradient material deposition has been exploited to construct an advanced injection molding tool which transitions from Invar in the center to stainless steel on the outside. The resulting tool exhibited minimal distortion from thermal stress and excellent exterior corrosion resistance.IntroductionA functional gradient material is a single, solid piece of material which exhibits spatially varying material properties. By creating such materials, one can tailor the composition of an artifact such that material properties are locally optimized. For example, for optimal tool life it is desirable to have a hard outside shell for wear resistance and a ductile core to resist brittle fracture. Traditionally, such benefits have been achieved through the coating or cladding of existing artifacts with shells of different physical characteristics.

Traditionally, such benefits have been achieved through the coating or cladding of existing artifacts with shells of different physical characteristics. While these techniques have been used extensively with great success, problems can result from sharp interfaces between dissimilar materials. Internal stresses and coefficient of thermal expansion mismatches can result in delamination, and the sharp interface can act as an initiation site for fracture. A monolithic structure with smoothly varying composition could alleviate some of these problems. Objects with extremely complex geometry can also be difficult to coat or clad effectively. If these complex parts were built in a layered fashion, any surface could be effectively coated regardless of the geometric complexity.Shape Deposition Manufacturing (SDM) [1, 2] presents the possibility to create metallic artifacts with functional gradient materials.

SDM uses laser deposition to fuse metallic powders onto a substrate and these powders can be mixed to create a range of alloys. For example, one can a produce a solid sample of stainless steel with a smooth variation in hardness from 20 and 45 Rockwell C (Re) simply by mixing hard and soft powders in the proper ratio. By depositing these materials in layers, SDM enables the production of objects in which functional gradient materials are used throughout the volume of the object. For example, artifacts with alternating layers of soft and hard materials are expected to be much less susceptible to failure through fatigue. As shown in bimaterial structures [3], much of the crack energy is dissipated in the soft layer, preventing the penetration of cracks into the harder layer. By using SDM, one could realize the benefits of these layered materials in objects of arbitrarily complex geometry.

By using SDM, one could realize the benefits of these layered materials in objects of arbitrarily complex geometry. This paper will examine some of the benefits and proPerties of the functional gradient materials produced with SDM. In particular it will examine the homogeneity of the deposits and the resulting material properties. Finally, it will present an example artifact which benefits from the use of functional gradient materials.521Laser Deposition SystemThe laser deposition system for producing functional gradient materials is essentially the same system described in Reference [2]. A 2400 W, continuous wave Neodymium:YAG laser is used to fuse metallic powders onto a substrate. The laser is focused onto the substrate where the transferred energy creates a melt pool. Metallic powders from three different powder feeders are fed from a single powder feed nozzle into the melt pool and subsequently melted.

Metallic powders from three different powder feeders are fed from a single powder feed nozzle into the melt pool and subsequently melted. Mixing of the powders occurs in an input funnel and along the length of a 1.5m feed tube. The feed rate of each of the powders and laser power are controlled automatically and the deposition apparatus is moved across the surface of the part using a four degree-of-freedom robotic manipulator. By using this the composition of the metal deposit at any point on the surface can be accurately system, controlled.Alloying of Metal PowdersThe fITst issue regarding the mixing of metallic powders for laser deposition is whether a homogeneous material is formed during solidification. Because of inadequate mixing or melting, the resulting material could possibly be comprised of isolated islands of one material in a matrix of the second material.

Because of inadequate mixing or melting, the resulting material could possibly be comprised of isolated islands of one material in a matrix of the second material. Such behavior would result in material properties significantly different than expected for a full alloy.To investigate the degree of alloying, electron backscatter microscopy was performed on samples of 50% Invar and 50% 316L stainless steel. Such a material combination is of interest for SDM because Invar exhibits a very small coefficient of thermal expansion and stainless steel has high corrosion resistance. A combination of the two results in a part which has minimal distortion due to thermal stresses and can withstand the acid etch required to remove the copper support structure. Figures 1 and 2 show the composition of chrome and nickel for samples A and B versus location. In sample A, a sharp interface between the two materials was achieved by depositing stainless steel next to an already existing deposit of Invar.

In sample A, a sharp interface between the two materials was achieved by depositing stainless steel next to an already existing deposit of Invar. The transition from one material to the other occurs over the length of 1.5mm, which is approximately the overlap of one laser pass on another. Sample B was deposited in a continuous fashion where the composition was varied between 100% Invar and 100% stainless steel in 10 discrete increments over 50mm. Electron backscatter images at the midpoint of each sample are shown in Figures 3 and 4 for samples A and B respectively. In electron backscatter images, elements with high average atomic number apPear lighter than those with low average atomic number. Figure 3 shows a distinctly lighter region of high atomic number material (Invar) separated from the darker, lower atomic number region (stainless steel) by a jagged interface. Figure 4 shows an image from sample B where the composition is 50% Invar and 50% stainless steel.

Figure 4 shows an image from sample B where the composition is 50% Invar and 50% stainless steel. This image has essentially uniform intensity, If the powders were not fully melting and mixing, one indicating a homogeneous composition. would eXPect to see pockets of light and dark on the order of the particle diameter (100 Jlm). Separate X-ray diffraction analysis shows that only a single phase is present in the material. This indicates that the constituents have formed a single-phase, solid-solution and are fully alloyed.The alloying properties of Invar and stainless steel should be contrasted with those of Aluminum-bronze and 316L stainless steel. Deposits which transition from 100% bronze to 100% stainless steel show visible segregation of the two materials into distinct bands and significant cracking. Figure 5 shows an electron backscatter image of such a composition where there are dark regions of stainless steel surrounded by the lighter bronze.

(Sample B)1040523Figure 3: Electron backscatter image for Sample A at location of 2mm. Lighter shades indicate higher atomic number of Invar.Figure 4: Electron backscatter image ofSample B at midpoint. Uniform shade indicates complete alloying of 50% Invar and 50% stainless steel.Figure 5: Electron backscatter image of aluminum-bronze and 316L stainless steel mixture. Note the distinct islands of stainless steel (dark) within the largely bronze matrix.524Properties of Graded MaterialsGiven that certain materials will fully alloy during the deposition process, one would expect the material properties in the graded sections to transition smoothly between the two extremes. This is indeed what occurs. For instance, in the Invar to stainless steel graded material illustrated in Figure 2, there is a continuous distribution of oxidation resistance which scales with the level of chrome.

For instance, in the Invar to stainless steel graded material illustrated in Figure 2, there is a continuous distribution of oxidation resistance which scales with the level of chrome. Invar, which has no chromium will oxidize to dark blue at elevated temperatures while stainless steel will remain silver. Figure 6 shows images of both the graded sample (Sample B) and the sharp transition samples (Sample A) after oxidation in a 400°C furnace for 25 minutes. Note the continuous distribution of oxidation in the graded sample.100% Invar100% 316L100% Invar100% 316LSample A: sharp transitionSample B: graded transitionFigure 6: Oxidation of sharp (Sample A) and graded transitions (Sample B) from Invar to stainless steel.Invar-stainless steel transitions are especially important for the SDM process because the presence of Invar has been found to drastically reduce the amount of deformation resulting from residual thermal stresses.

Pure Invar deposits show approximately 1/2 the deflection of pure stainless steel deposits of the same geometry [2]. Laser-deposited Invar, however, cannot withstand the long term immersion in nitric acid that is required to remove the copper support material used in the fabrication of metal parts. A thin layer of stainless steel is sufficient to protect the part in the acid bath, and beams made with a core of Invar and a thin shell of stainless steel have been found to have only 2/3 the deflection of pure stainless steel beams.As stated in the introduction, one of the most useful functional gradient materials involves a hard outer shell for wear resistance surrounding a more ductile core. The laser deposition system has been used to produce samples which continuously transition from 316L stainless steel (Re < 20) to a harder, corrosion resistant alloy (Re =43) similar to 414 stainless steel.

Figures 7 and 8 show the distribution of hardness and iron content versus location for a sample of this graded material which transitions from 100% hard alloy to 100% 316L stainless steel over 50mm. Note that while there is essentially a linear distribution of the iron content, the hardness distribution is distinctly non-linear. Most of the hardness variation occurs between 20 and 30mm which corresponds to concentrations of 40 to 60%. The plateaus on either side of this transition region correspond closely to the hardness of the major constituent. that hardness on the Rockwell C scale does not scale linearly with carbon in stainless steels and the nonlinearity can be qualitatively explained by the variation of carbon content.

This tool, which is shown schematically in F~gure 9, is comprised of three materials: Invar, stainless steel and copper. The bulk of the tool IS made from Invar to reduce deformation due to residual thermal stresses. A thin shell of stainless steel surrounds the part to protect it during the acid etch process. To minimize cycle time, cooling channels would ideally be placed directly under the mold section of the tool; this, however, would not allow for ejector pins for part removal. As a compromise, solid deposits of copper have been located 2mm below the mold surface to facilitate heat removal form the part during production. Conformal cooling channels are located 2mm on either side of the copper to facilitate heat removal from the copper.

Conformal cooling channels are located 2mm on either side of the copper to facilitate heat removal from the copper. The proximity of the copper to both the cooling channels and the plastic part should significantly decrease the cycle time for each part by reducing the time necessary for temperature stabilization.Near net shape deposition of each layer began with Invar at the center of the part and spiraling outward following the geometry of that layer. In the outer few passes, a con~inuous transition was made from Invar to stainless steel to assure all outside surfaces were corrosion Internal features such as the cooling channels and the pocket for the copper deposit, resistant. were machined layer-by-layer as the part was deposited, while outside features were machined and electro discharge machined (EDM) after the entire part was deposited. Figure 10 shows a cross section of the part which shows the relative locations of Invar, stainless steel and copper.

Figure 10 shows a cross section of the part which shows the relative locations of Invar, stainless steel and copper. Construction of this part illustrated that continuous transitions between materials could be made within a single deposited layer. Future work will include tools with very hard outside shells to decrease tool wear and Invar cores to reduce thermal distortion.ConclusionsThe ability to specify material properties at any location within an artifact presents numerous benefits to designers in terms of choosing the optimal material for each application. Recently, the laser deposition system within SDM has been modified to enable the continuous deposition of graded metals from up to three constituent feed powders. This, coupled with the inherent flexibility of SDM, now enables the production of complex three dimensional artifacts with continuously varying material properties throughout the volume of the artifact.

Materials which traditionally alloy also alloy with this system to form homogeneous, single phase deposited material. In general these graded materials exhibit varying material properties with varying composition but the property variations are not necessarily linear with concentration. The new system has been used to produce samples with continuously variable oxidation resistance and hardness. Future work on the material properties of these graded materials will focus on their fracture and fatigue properties as compared to sharp material interfaces.The system has also been utilized to fabricate an advanced, multimaterial injection molding tool. The tool has copper deposits and cooling channels to minimize cycle time and Invar core to reduce distortion due to residual thermal stresses. The outside of the tool is stainless steel to prevent corrosion while in the acid bath used to remove the copper used as a sacrificial support material.

The outside of the tool is stainless steel to prevent corrosion while in the acid bath used to remove the copper used as a sacrificial support material. The Invar and stainless steel were deposited in a continuous, outward spiral, where the transition to stainless steel was made within the outer few passes along the contour. The same technique could be also used to produce long-life, wear-resistant tools with a hard, corrosion resistant outer shell. This technique can be extended to any multimaterial application where stress concentrations due to sharp material interfaces may limit performance.527Figure 9: Model of advanced ALCOA injection molding tool.

Tool is made of Invar, stainless steel and copper and has two conformal cooling channels in each half to remove heat quickly from the part.Figure 10: Horizontal cross-section of one half of advanced injection molding tool illustrating gradient from Invar to stainless steel on surface of part.AcknowledgmentsThe authors gratefully acknowledge the Defense Advance Research Projects Agency (#F33615-94 C-4450 and #F33615-95-C-5560), the Office of Naval Research (#NOOOI4-96-1-0354), the National Science Foundation (#MIP-9618050) and ALCOA for their support of this research.References[1] Merz, R., Prinz, F., Ramaswami, K., Terk, M., and Weiss, L. (1994) "Shape Deposition Manufacturing," Proceedings of Solid Freeform Fabrication Symposium, The University of Texas at Austin, pp. 1-8.[2] Metals for Shape Deposition Manufacturing," Proceedings of Solid Freeform Fabrication Symposium, The University of Texas at Austin, pp.

There are potentiallymany applications, where 3D micro-features can benefit the overall function of the macrostructure.One such application has been recently identified in cataract surgery where a medical device, called acoaxial phacoemulsifier, is inserted into the eye through a relatively small incision and used to breakthe lens apart while removing the lens pieces and associated fluid from the eye through a small tube.In order to maintain the eye at a constant pressure, the phacoemulsifier also includes an irrigationsolution that is injected into the eye during the procedure. It has been reported, however, that theimpinging flow from the irrigation solution on the corneal endothelial cells in the inner eye damagesthese cells during the procedure.

It has been reported, however, that theimpinging flow from the irrigation solution on the corneal endothelial cells in the inner eye damagesthese cells during the procedure. As a result, we are exploring methods for reducing flow velocitiesduring this procedure, and have designed a complex, 3D micro-vane within a sleeve that introducesswirl into the irrigation solution, and thus, produces a flow with rapidly dissipating flow velocities.However, the fabrication of the sleeve could not be accomplished using µSL alone, and thus, a two-part design was accomplished where a sleeve with the micro-vane was fabricated with µSL and athreaded fitting used to attach the sleeve to the phacoemulsifier was fabricated using an Objet Eden333 rapid prototyping machine. The new combined device was tested in a water container usingparticle image velocimetry, and the results showed an ejection of the irrigation fluid through the micro-vane in three different radial direction.

It is believed that this new device will reduce damage toendothelial cells during cataract surgery and significantly improve patient outcomes from this procedure.This unique application demonstrates the utility of combining µSL with a macro rapid prototypingtechnology for fabricating a real macro-scale device with functional, 3D micro-scale features that couldnot be fabricated using conventional methods.Keywords: Phacoemulsification, Cataract Surgery, Irrigation, Swirl, Microstereolithography5531. IntroductionIn cataract surgery, phacoemulsification has been widely used to remove the crystalline, where anultrasound irrigation and aspiration hand piece emulsifies and removes the internal lens of the eyethroughincision sites, and an artificiallensisinsertedimmediately.There aretwotypes ofphacoemulsification used in the cataract surgery: bimanual and coaxial.

Bimanual phacoemulsificationemploys two separate stainless pipes for irrigation and aspiration, whereas coaxial phacoemulsificationemploys a combination of stainless pipe named ‘tip’ and a dead-ended plastic pipe named ‘sleeve’,which coaxially surrounds the tip, and where a gap between the tip and sleeve leads irrigation flow.Generally, bimanual phacoemulsification has the advantage of a small incision (1.2 mm – 1.4 mm) incomparison to a relatively large incision (2.8 mm – 3.2 mm) in coaxial phacoemulsification (Gajjar, etal., 2007; Osher and Injev, 2007). Contrarily, coaxial phacoemulsification has advantages of highvacuum level, better temperature control, integrity of the incision, easy handling, less invasive and saferoperation, better sealability, and fluidic benefits (Osher and Injev, 2007).

Even though the two typesof phacoemulsification have advantages and are still in use, it has been reported that the bimanualphacoemulsification has more disadvantages such as limited irrigation by small incisions (Osher andInjev, 2007), chamber instability due to leakage through two incision sites (Lee et al., 2009), greateringress of bacterial inoculums (Gajjar et al., 2007), and development of mechanical wound trauma dueto angular movement of the instruments through the small incision sites (Lee et al., 2009).

In therecent advancement, microcoaxial phacoemulsification with an incision of 2.2 mm has been developedsuch that it can overcome the drawbacks of bimanual and the problem of the large incision of coaxialphacoemulsification, as still having the advantages of coaxial phacoemulsification mentioned above(Gajjar et al., 2007, Lee et al., 2009).The irrigation solution in phacoemulsification surgery plays an important role in maintaining aconstant pressure in the anterior chamber and dissipating heat generated from the ultrasonic tip as wellas maintaining an endothelial cell survival by the chemical composition of the solution (Joussen et al.,2000).

Despite of the importance of the irrigation solution, several studies have shown that theimpinging flow from the irrigation solution on the corneal endothelial cells in the inner eye damagesthese cells during the procedure (Polack and Sugar, 1976; Binder et al., 1976; Miller et al., 1996;Hayashi et al., 1996; Topaz et al., 2002; O’Brien et al., 2004; Steinert and Schafer, 2006). It is wellknown that cells including the one in the cornea endothelial layer can be damaged by the tangentialcomponent of the hydrodynamic forces, i.e. shear stress, resulting from irrigation (Binder et al., 1976;Born et al., 1992; Thoumine et al., 1995; Kaji et al., 2005).As a result ofthe effort to reduce flow velocities during this procedure, it has beensuggested to design and fabricate a complex, 3D micro-vane within a sleeve that introduces swirl intothe irrigation solution, and thus, produces a flow with rapidly dissipating flow velocities.

Althoughadditive manufacturing (AM) systems are capable of producing 3D complex structures, the fabricationof the sleeve with 3D micro-vanes could not be accomplished in one AM system alone because the554sleeve with the micro-vane has both macro- and micro-features.

That is, most of commercialized AMsystems cannot meet resolution for micro-features, whereas there are several types of lab-based, additivemicrofabrication systems, which are not suitable for producing macro-structures because oflongfabrication time, distortion, weak mechanical strength, and limitation of work volume (Varadan et al.,2001; Choi et al., 2006; Regenfuss et al., 2007; Khalil and Sun et al., 2007; Lee et al., 2008).Therefore, a two-part design is needed, where a sleeve with the micro-vane can be fabricated with amicrofabrication system and a threaded fitting used to attach the sleeve to the phacoemulsifier can befabricated using a commercialized AM system so that they can be assembled after fabrication.

Among3D microfabrication systems for producing a sleeve, it is thought that microstereolithography (µSL), atechnologythat evolved from stereolithography (SL),is more compatiblethan others duetoitscapability of achieving high resolution (< 10 µm), and high-aspect ratio fabrication (Bertsch et al.,1997; Sun et al., 2005; Choi et al., 2009a, 2009b). In addition to the microfabrication, commercialAM systems with high-resolution are recommended for producing a threaded fitting part.In this work, a two-part design was accomplished, and then a 3D micro-vane within a sleeveand a threaded fitting were fabricated with µSL and a commercial AM system, respectively. Comparedto a commercial AM system, µSL has more issues associated with microfabrication, for example,requirement of low viscosity material, cure depth control, dimensional accuracy, and distortion afterbuilding.

For the successful fabrication of a sleeve, the work presented here includes system alignmentusing a large imager, preparation of a low viscosity photocurable resin using a diluent, cure depthcontrol using a light absorber, curing characteristics test, and multiple exposure for obtaining a micro-vane within a sleeve without bending.Finallythetwofinal parts wereintegrated with aphacoemulsifier, and particle image velocimetry (PIV) was used to show the reduction of mean velocitydistributionina watercontainer.Thefollowings willdescribethetwo-partdesign,micro/macrofabrication, and PIV results in more detail.2. Experimental apparatus2.1 Multi-Jet Modeling (MJM)MJM, also referred to as 3D Printing (3DP), is one of the Additive Manufacturing (AM) processesavailable, which uses a printer head, ultraviolet light (UV), and a photocurable liquid resin.

In thiswork, an Objet Eden 333 (Objet Geometries Ltd., Billerica, MA) was used for producing a threadedfitting. This system is capable of stacking thin layers with a 16 µm thickness by injecting a curingviscous photopolymer through the printer head while simultaneously curing the photopolymer with theUV light. In particular, it utilizes a build and support material separately, and after finishing a build,the support material is removed by using high-pressure water.2.2 Dynamic mask microstereolithography (µSL)For the purpose of producing 3D microstructures, µSL has been widely used for various engineeringapplications since its first introduction (Ikuta and Kirowatari, 1993).

µSL technology is based on555stereolithography (SL), and it can be classified into two types: vector-by-vector µSL, also referred toas scanning µSL, by scanning a focused laser beam (Ikuta and Kirowatari, 1993; Zissi et al., 1996;Sun and Zhang, 2002; Lee et al., 2007, 2008); and integral µSL, also referred to as projection µSL,by projecting an image reflected on a dynamic mask (Bertsch et al., 1997, 2001, 2004; Sun et al.,2005; Choi et al., 2006, 2009a, 2009b; Limaye and Rosen, 2007; Ha et al., 2008; Han et al., 2008).In this work, Digital Micromirror Device (DMDTM)-based µSL was used to produce a sleevewith a micro-vane. Fig.

Fig. 1 shows the µSL system, which consists of an Omnicure S2000TM as a lightsource with an output of 200 W, a collimating lens, and an optical fiber (EXFO Co., Canada), aDMDTM with ~780,000 micromirrors with the size of 13.68 µm (DMD Starter Kit, Texas Instruments,USA), a tube lens with a focal length of 120 mm (Achromat doublet lens, MellesGriot Co. USA), analuminum-coated reflecting mirror, a focusing unit (IM-4, Nikon Co., Japan), an objective lens (CFIPlan Flour 4×, Nikon Co., USA), a motor-driven linear Z-stage with a resolution of 1 µm (ATS100-050, Aerotech, USA), and a stainless resin vat. For the fabrication of microstructures, a desired 3Dmodel is sliced and converted into binary images as cross-sections, which are transferred to theDMDTM board in process. The DMDTM plays a key role in dynamic mask projection µSL, and eachindividual micromirror is tilted at ±12° according to the binary value, i.e., 0 or 1, so that entireDMDTM forms a certain image.

The DMDTM plays a key role in dynamic mask projection µSL, and eachindividual micromirror is tilted at ±12° according to the binary value, i.e., 0 or 1, so that entireDMDTM forms a certain image. The illuminated light is patterned by being reflected on the DMDTM,delivered to the resin surface through several optics, and finally the focused pattern cures the resinsurface with the prescribed thickness. Finally, the Z-stage stacks the cured layers and refreshes newresin surface for the next build, and this continuous builds produce a 3D microstructure. In thedeveloped system, the exposure energy at the resin surface is ~26.35 mJ/cm2, and the reduction ratio is~0.329 between an image on the DMDTM and resin surface. Although the µSL system is describedbriefly here, the interested readers are recommended to read the references (Choi et al., 2006, 2009a,2009b) for further details of the system.Fig.

Although the µSL system is describedbriefly here, the interested readers are recommended to read the references (Choi et al., 2006, 2009a,2009b) for further details of the system.Fig. 1 Developed µSL system: (a) schematic of the system, (b) photo of the system3. Materials, Design, and Methods5563.1 MaterialsObjet Eden 333 system using Objet FullCure® 840 as a build material and FullCure® 705 as a supportmaterial was utilized to fabricated the threaded fitting. ProtoThermTM 12120 was used as a mainmaterial in µSL (DSM Somos®, New Castle, DE, USA).

ProtoThermTM 12120 was used as a mainmaterial in µSL (DSM Somos®, New Castle, DE, USA). Propoxylated (2) neopentyl glycol diacrylate(PNGD) (Sartomer Inc., Warrington, PA) was used.To control cure depth, Tinuvin 400 (CibaSpecialty Chemicals, Newport, DE) was added to the ProtoTherm-PNGD mixture.3.2 Design of sleeveTo replace the original sleeve in a phacoemulsifier with a micro-vane within a sleeve, a two-partdesign for a threaded fitting fabricated in MJM and the sleeve with the micro-vane fabricated in µSLwas accomplished. The threaded fitting as shown in Fig. 2 consists of a thread, inner channel, andspace for the sleeve to be equipped. The sleeve design has a micro-vane as shown in Fig. 3, and thedimensions referred an existing sleeve and tip. The micro-vane as shown in Fig.

The sleeve design has a micro-vane as shown in Fig. 3, and thedimensions referred an existing sleeve and tip. The micro-vane as shown in Fig. 3 was generated bysweeping a rectangular section with the size of 0.219 mm × 0.110 mm along with a spiral curve.The spiral curve consists of two curves, which have a diameter of 1.54 mm and different pitch andrevolution as shown in Fig. 3 (b). The upper curve has 1 mm height and 2 mm pitch, and the lowercurve has 2 mm height and 10 mm pitch, so that the combined curve has steep and gentle slopesfrom the bottom to top, which introduces smooth swirl. Fig. 4 shows the 3D model of the assembledsleeve with the threaded fitting.Fig. 2 3D model of the threaded fitting (unit: metric)557(a) (b)Fig. 3 3D model of the sleeve with micro-vane: (a) hollow tube and section, (b) micro-vane (unit:metric)(a) (b)Fig.

2 3D model of the threaded fitting (unit: metric)557(a) (b)Fig. 3 3D model of the sleeve with micro-vane: (a) hollow tube and section, (b) micro-vane (unit:metric)(a) (b)Fig. 4 3D model of the assembled sleeve: (a) 3D model, (b) section3.3 Methods3.3.1 Uniform light intensityTo obtain uniform light intensity on the resin surface, where non-uniform intensity causes unevenphotopolymerization and distortion of the final microstructure, µSL was manually aligned and theintensity profile was measured using a large area imager (LBA-USB-L11058, Ophir-Spiricon Inc.,Logan, UT). Two cross-sections of the design of sleeve were used to generate patterns on the imager.3.3.2 Viscosity controlGenerally, µSL requires the use of a low viscosity photocurable resin, because it uses a deep-dipprocess to refresh the resin surface (Varadan et al., 2001; Choi et al., 2009a).

The main material,DSM Somos® ProtoThermTM 12120 as a commercial resin, has ~410 cP as a nominal viscosity at 30558oC, which is relatively high to be used in µSL. To reduce the viscosity, ProtoThermTM was blendedwith PNGD as a diluent, and the viscosity was measured using a Brookfield viscometer (DV-E,Middleboro, MA) according to the mixture ratio such as 100: 0, 90: 10, 80: 20, 70: 30, 60: 40, and50: 50 (ProtoThermTM: PNGD) at room temperature (24 oC). In µSL, viscosity values small than 200cP is recommended for fast recoating with thin layer thickness (Varadan et al., 2001; Choi et al.,2009a, 2009b)3.3.3 Cure depth controlIn general, a commercial resin is compatible with macro-scale structures, and it has relatively largepenetration depth of the light (Dp), e.g., 0.14 mm and low critical exposure energy (Ec), e.g., 12.2 mJ/cm2 for ProtoThermTM.

Large Dp leads overcure and occlusion in overhanging micro-features, thuscure depth needs to be controlled (Choi et al., 2009a). In addition to the large Dp, the main materialwas diluted, thus there exists no information about curing characteristics (Dp and Ec), which would beused for determining process parameters such as exposure energy and others. First,, to control curedepth, Tinuvin 400 as a light absorber at a concentration of 1.0 wt% was used, and in order todetermine the specific curing characteristics of the mixture a curing experiment was conducted. Dpand Ec can be calculated from the Beer-Lambert formula provided in Eq. (1) (Jacobs, 1993). Curedepth (Cd) was measured using a crossbeam, which was produced by varying exposure energy (Emax) atthe resin surface (Zissi et al., 1996; Choi et al., 2009a, 2009b). The exposure energy was controlledby adjusting exposure time of the lamp such as 10 s, 12 s, 14 s, 16 s, 18 s, 20 s, and 22 s.

The exposure energy was controlledby adjusting exposure time of the lamp such as 10 s, 12 s, 14 s, 16 s, 18 s, 20 s, and 22 s. Eightcrossbeams were built for each exposure time.Cd = Dp ln (Emax/Ec)(1)3.3.4 Multiple exposure for segmented regionEven though cure depth can be controlled using a light absorber, a desired microstructure havingoverhanging micro-features with different slopes may not be obtained (Choi et al., 2009a). This isbecause an overhanging feature having a gentle slope is more affected by the penetration depth than afeature having a steep slope. Contrarily certain amount of overcure is important to ensure that thecurrently cured layer is attached to the previously built layer, and weak exposure energy often leads todistortion of microstructures after building. Thus, both of cure depth control and overcure play animportant role in producing microstructures.

Thus, both of cure depth control and overcure play animportant role in producing microstructures. However, both of them cannot be satisfied at the sametime because cure depth control directly affects an amount of overcure.. This contradiction may bepartially solved by process planning according to a geometry (Limaye and Rosen, 2006, 2007) andmultiple exposures for segmented regions (Park et al., 2009). In this paper, the micro-vane and hollowtube features as shown in Fig. 3 were separately cured using multiple exposures with different energy.This was accomplished by saving and slicing separate solid models for the micro-vane and hollow tubeand using different exposure times for each model.

That is, a relatively long exposure time was used559for the hollow tube to ensure that it can stand alone and any distortion does not occur after build,whereas a short exposure time was used for obtaining fine features of the micro-vane, which does notpreventthe flow of solution.By comparing micro-vanes fabricated with and without multipleexposures for segmented regions, the introduced method was verified. Sleeves with the micro-vanewere fabricated using the process parameters as described in Table 1.

Sleeves with the micro-vanewere fabricated using the process parameters as described in Table 1. The rectangular section ofmicro-vane with the size of 0.219 mm × 0.110 mm was measured using a stereomicroscopic (LeicaMZ16, Leica Microsystems, Switzerland) to confirm the feasibility of the use of the multiple exposure.Table 1 Process parameters for producing a sleeve in µSLTotal layer number832Layer thickness (µm)18Settling time (sec)10Irradiance (mW/cm2)26.4Single exposure time (sec)15 ~ 20Multiple exposure time (sec)Sleeve25Vane15 ~ 20Total fabrication time (hour)7 ~ 103.3.5 Flow velocimetryUsing the fabricated sleeve with the micro-vane, flow velocimetry was measured through PIV.

Tracerparticles (S-HGS Silver-coated hollow glass sphere, Dantec dynamics ) were mixed in the irrigationfluid and water container, and pulsed laser light sheet, which was produced by Nd-YAG laser (30mJ/pulse, 15 Hz, Solo PIV, New Wave Research Inc), illuminated the particles. The laser light sheetwas radiated twice within a short interval of time, and a CCD camera (TM-1325CL, JAI/Pulnix Inc)capturedthese scatteringlightimages ontwo successive frames.By usingtheseimages,displacement of the tracer particles in the image was analyzed by cross-correlation function. Toevaluate the physical position and displacement of tracer particles in the image, a calibration plate,which has grid holes with 0.5 mm intervals, was inserted into the measurement plane in a fluidimaged by the CCD cameras. Based on this image, the pixel position named ‘Image coordinate’ wasassociated with the physical position by the second order polynomial warping function.

Based on this image, the pixel position named ‘Image coordinate’ wasassociated with the physical position by the second order polynomial warping function. The velocityof the tracer particles was obtained by dividing the displacement of the tracer particles by interval timeof the radiated laser light sheets.4. Results and Discussion4.1 Uniform light intensityThe measured intensity profiles as shown in Fig. 5 show that the focused pattern on the resin surfacehas a uniform intensity, thus it can be thought that the system was properly aligned for achievingmicrostructures. In particular, because the designed sleeve has both the thin wall and fine micro-vane,a non-uniform intensity on the resin surface cause a distortion of the final sleeve. Additionally,560thebending ofthe sleeve might occur becausethe sleeveis arelative high-aspectratiomicrostructure.

Additionally,560thebending ofthe sleeve might occur becausethe sleeveis arelative high-aspectratiomicrostructure. However, it is noted that distortion and bending of the sleeve may also be affected byexposure energy and other process parameters. Thus, it could be expected that the aligned systemwith the uniform intensity would take a share of producing a straight sleeve without distortion.(a)(b)Fig. 5 Light intensity profile: (a) bottom section. (b) top section with vane (Note that the color barsin the figures represent not absolute values, but relative values. The average intensity of the greenpart was ~26.35 mJ/cm2 ).4.2 Viscosity controlFig. 6 shows the measured viscosity of the ProtoThermTM 12120 dialuted with PNGD. The graphshows that the viscosity exponentially decreased as the amount of PNGD increased.

6 shows the measured viscosity of the ProtoThermTM 12120 dialuted with PNGD. The graphshows that the viscosity exponentially decreased as the amount of PNGD increased. The initiallymeasured viscosity without the diluents at room temperature was ~1100 cP, which is relatively higherthan the viscosity (~410 cP) provided from the company, and it is thought that the viscosity increasedas time elapsed. The viscosity for a 50:50 mixture (ProtoThermTM: PNGD wt%) was measured to be~100 cP, and this mixture was chosen for producing the sleeve.Fig. 6 Viscosity variation according to the mixture ratio561(~7.4)Fig. 7 Measured cure depth by varying exposure energy (A natural log scale was used for X axis andthe error bars represent plus/minus one standard deviation).4.3 Cure depth controlCure depth was controlled by blending Tinuvin 400 as a light absorber, and Fig. 7 represents measuredcure depths according to energy dosage on the resin surface.

7 represents measuredcure depths according to energy dosage on the resin surface. By regressing the measured data, Dp and Ec were determined ~108 µm and 145 mJ/cm2. the curing characteristics about ProtoThermTM as a main material, Dp was decreased ~22.9 % and Ec was increased ~1089 %. Even though the calculated Dp is small than the original value (~140 µm) of ProtoThermTM, it is relativelyComparedtolarger than that used by other researches (Choi et al., 2009a, 2009b). However it was expected that aminimum achievable feature size would be about ~ 100 µm, and it is a relevant value for producingmicro-vane.4.4 Multiple exposureThe thickness and width of micro-vane according to the exposure time were measured and Fig. 8shows the results for multiple and single exposure. The measured micro-vane thickness using multipleexposure was 40 % ~ 81 % larger than the designed dimension (110 µm), whereas the thickness usingsingle exposure was 121 % ~ 319 % increased.

The measured micro-vane thickness using multipleexposure was 40 % ~ 81 % larger than the designed dimension (110 µm), whereas the thickness usingsingle exposure was 121 % ~ 319 % increased. Additionally, micro-vane widths using multiple andsingle exposure were 38 % ~ 40 % and 21 % ~ 45 %, respectively, smaller than the designed width(219 µm). Using multiple exposure it was possible to obtain thinner micro-vane, it was, however,detached or broken from the sleeve. Even though the exact dimensions were not achieved, multipleexposure fabrication has better result of obtaining thinner micro-vane. Fig. 9 shows the fabricatedmicro-vane inside the sleeve with and without multiple exposure when the exposure time was 15 s.562According to Fig. 9 it might occur that the overcured micro-vane by single exposure occludes the pathof solution flow. Additionally, relatively short exposure time often resulted in distortion in the sleeveas shown in Fig.

Additionally, relatively short exposure time often resulted in distortion in the sleeveas shown in Fig. 10 when rinsing and post-curing. Thus it can be thought that multiple exposure iseffective mean to obtain fine feature of the micro-vane. Finally it is recommended that exposure timesfor sleeve and micro-vane have to be more than 20 s and 15 s ~ 20 s, respectively. Fig. 11 showsthe fabricated sleeve using exposure time of 30 s and 20 s for the hollow tube and micro-vane.(a) (b)Fig. 8 Measured micro-vane dimensions with and without multiple exposure: (a) micro-vane thickness,(b) micro-vane width(a) (b)Fig. 9 Optical micrographs of micro-vanes fabricated using 15 s exposure time: (a) with and (b)without multiple fabrication.563Fig. 10 Optical micrographs of distorted sleeveFig.

9 Optical micrographs of micro-vanes fabricated using 15 s exposure time: (a) with and (b)without multiple fabrication.563Fig. 10 Optical micrographs of distorted sleeveFig. 11 Successfully built sleeve using multiple exposure4.5 IntegrationThe sleeve with micro-vane fabricated with the µSL system and the threaded fitting fabricated with theEden 333 system were integrated as shown in Fig. 12. Both of sleeve and threaded fitting weresuccessfully built for integration and Fig. 13 shows a phacoemulsifier with the fabricated sleeve andthreaded fitting.564Fig. 12 Integrated sleeve and threaded fitting.(a)(b)Fig. 13 Phacoemulsifier equipped with 3D micro-vane within the sleeve: (a) phacoemulsifier with thefabricated sleeve, (b) magnified image of the sleeve4.6 Flow velocimetry565We measured the velocity distribution with 0.5mm intervals when the tip was submerged in the watercontainer. Measurement plane is normal to the axis of the tip.

Measurement plane is normal to the axis of the tip. Fig 14 shows mean velocity vectors.The position of the measurement planes located at 1.0mm, 1.5mm, 2.0mm, 2.5mm and 3.0mm fromthe head of the tip. Color indicates magnitude of velocity. It can be confirmed that water wasirrigated from the head of tip, and irrigation gushed in three radial directions.Fig. 14 Flow velocimetry.5. ConclusionsIn this work, the successful fabrication in use of sleeve with a 3D micro-vane, which is used inphacoemulsification for cataract surgery, was demonstrated. The 3D micro-vane was suggested forintroducing swirl into the irrigation solution and dissipating flow velocities from the irrigation solutionon the corneal endothelial cells in the inner eye. To fabricate 3D micro-vane within a sleeve, a two-part design for a threaded fitting, which can be assembled with a phacoemulsifier, and sleeve withmicro-vane, which can be integrated with the threaded fitting.

Both of designed parts were fabricatedusing a commercial additive manufacturing system and microstereolithography system. Compared tofabrication in commercial additive manufacturing system, microfabrication required reforming of materialpropertiesandadjusting process parametersfor producing micro-features.Toaccomplishmicrofabrication in microstereolithography, the fabrication system was manually aligned and verified byinspecting pattern intensity, and it was confirmed that the focused light had uniform light intensity.Cure depth of the material were controlled by blending a light absorber, and finally they were suitablefor producing 3D micro-vane. Additionally, for obtaining a non distorted sleeve with 3D micro-vane,multiple exposure was conducted, and results show that the suggested multiple exposure is a promisingmean for producing a final sleeve with 3D micro-vane.

Finally the new combined device was testedin a water container using particle image velocimetry, and the results showed the irrigation fluids566ejecting through the micro-vane towards the three radial directions. In conclusion, it is believed thatthis new device will reduce damage to endothelial cells during cataract surgery and significantlyimprove patient outcomes from this procedure, and this unique application demonstrates the utility ofcombining µSL with a macro rapid prototyping technology for fabricating a real macro-scale devicewith functional, 3D micro-scale features that could not be fabricated using conventional methods. Inthe future, the phacoemulsifier with the developed sleeve will be tested in cow eyeballs to showsignificant reductions in flow velocities and resulting shear stresses on the endothelial cells.ReferencesA. Bertsch, S. Zissi, J.Y. Jézéquel, S. Corbel, and J.C. André, Microsyst. Technol. 3, 42-47 (1997).A. Bertsch, S. Zissi, J.Y.

nctionally Graded Copper – Steel Using Laser Engineered Net Shaping  ProcessFredrick F. Noecker II and John N. DuPontDepartment of Materials Science and Engineering, Lehigh University Bethlehem, PA 18015AbstractLaser Engineered Net Shaping (LENS) is an emerging Solid Freeform Fabrication(SFF) process capable of producing fully dense metallic parts with complex shapes directly from a computer-aided drawing (CAD) without the need for molding or tooling. The LENS process also shows promise in producing components with graded compositions. One potential application is the production of steel – copper die casting materials. Copper is currently deposited on dies made out of steel to enhance thermal management, however, difficulties can occur at the Fe-Cu interface, such as, a large solidification temperature range in the Fe-Cu system (which can cause solidification cracking), formation of undesirable phases, and differences in coefficient of thermal expansion.

The current research goal is to develop LENS processing parameters to optimize the deposition of functionally graded steel - Copper alloy. Dilution control experiments were carried out to determine the cracking susceptibility of Steel - Cu alloys of varying compositions. The resulting microstructures were characterized by various microscopy techniques. The influence of composition on the resultant cracking susceptibility and microstructure will be discussed. This information is useful for successful deposition of crack-free copper layers onto steel for die casting applications.(189 words in abstract)IntroductionLaser Engineered Net Shaping (LENS) is a growing solid free form fabrication process capable of producing fully dense 3-D complex shapes directly from a Computer Aided Design ( CAD) drawing via Direct Metal Deposition (DMD). LENS utilizes an Nd-YAG laser to produce a melt pool on a substrate attached to an X-Y table.

LENS utilizes an Nd-YAG laser to produce a melt pool on a substrate attached to an X-Y table. Powder metal from coaxial powder flow nozzles is injected into the melt pool as the table is moved along a pre-designed tool path thus producing a fully dense part by depositing line after line which become sequential layers. LENS has already proven its ability to produce molds with out of AISI H-13 Tool Steel (ref Mazumder 97) and shown promise in producing conformal cooling channels in molds made out of the same material (ref Knights 01). Additionally, several researchers have investigated the fabrication of functionally graded materials using direct metal deposition variant processes (ref Brice 00, Fessler 97, Li 99).AISI H-13 tool steel is a widely used mold material due to its favorable mechanical properties, but the thermal conductivity limits the melt cooling rate and subsequent mold cycle time.

The thermal conductivity of copper is nearly twice that of AISI H-13 tool steel at operating temperatures between 220-600 C (ref CarTech H-13 physical properties sheet, Touloukian 70).231Attempts have been made to clad materials with higher thermal conductivity, such as copper, to molds. This has resulted in limited success due to the difference in thermal coefficient of expansion between copper and steel, which creates larger thermal stresses at the clad interface ultimately resulting in failure. The tool and die industry would like to exploit the combined LENS capability of producing functionally graded materials and with conformable cooling channels to improve die thermal management which will subsequently increase productivity. By increasing productivity 20%, every fifth day of production is free, along with cascading manufacturing variables that can result in a marked increase in profits (ref.

By increasing productivity 20%, every fifth day of production is free, along with cascading manufacturing variables that can result in a marked increase in profits (ref. Beard 01)Proof of Concept ExperimentPure Cun o i t a r t n e c n o C u C g n i s a e r c n IThe LENS was used to create a Functionally Graded Material (FGM) cube of H-13 tool steel and copper to evaluate metallurgical factors. Copper was continuously graded into H-13 tool steel that was deposited on a plain carbon steel substrate. Table 1 contains alloy compositional information. Figure 1 is an as polished light optical micrograph of the longitudinal cross-section revealing cracking and porosity. Porosity was observed in the FGM as the concentration of copper increased. The porosity is believed to be a function of the copper powder’s dendrite morphology.

Porosity was observed in the FGM as the concentration of copper increased. The porosity is believed to be a function of the copper powder’s dendrite morphology. A spherical powder shape should provide the solution to the porosity problem.Tool SteelFigure 1: FGM montage of H-13 – Copper LENS depositTable 1 – Composition of alloys used in steel – copper FGM and dilution control experimentsCracking was also observed in the FGM in the region where copper gradient began. Copper has been shown by several researchers to promote solidification cracking/hot cracking in steel (ref Asnis 65, Vainerman 68, Cooper 74, Dixon 81). The susceptibility of steel – copper alloys to solidification cracking is evident in the Fe-Cu phase diagram, Figure 2 (Swartzendruber 90) which reveals a large solidification temperature range and very limited solid solubility of either element which will lead to large amounts of terminal liquid.

Additionally, the dihedral angle for copper – mild steel is below 25° throughout the solidification temperature range, thus promoting austentite grain boundary wetting by copper (ref. Salter 66).232Dilution Control Experimental ProcedureTo avoid crack susceptible steel – copper concentrations, the range of copper concentration in steel that promote cracking must first be known. Dilution control experiments (ref DuPont 96) were undertaken to determine this compositional cracking range. Dilution experiments utilizing a Gas Tungsten Arc Welder (GTAW) with an automated cold wire feed provided an elegant way to easily produce a wide range of steel – copper compositions, knowing the filler metal and base metal compositions (see Table 1). Copper bead on rolled 1” wide x ¼” thick AISI 1013 steel rolled bar welds were made with all weld processing parameters constant except wire feed speed.

Copper bead on rolled 1” wide x ¼” thick AISI 1013 steel rolled bar welds were made with all weld processing parameters constant except wire feed speed. AISI 1013 steel was selected as a simplified, yet fairly representative system for AISI H-13 tool steel. Also, it is hoped that the results produced by welding copper onto steel under a condition of low restraint will be used in a broad range of applications beyond DMD.Figure 2: Fe-Cu Phase DiagramThe GTAW was set up with a .01” arc gap, 2mm/s travel speed, 250 amp current and 10 volt potential. The shielding gas was commercially pure Ar. The steel rolled bar was held along its entire length by a welding jig until the sample cooled to the touch. Volumetric filler metal feed rate is a function of wire feed speed, which was varied to alter the composition of copper in the fusion zone.

Volumetric filler metal feed rate is a function of wire feed speed, which was varied to alter the composition of copper in the fusion zone. Samples were fabricated with compositions ranging from AISI 1013 Steel 7 wt% Cu to 98 wt% Cu.Weld transverse cross sections of each welding condition were sectioned, mounted, ground and polished using standard Metallographic techniques, then etched in 2% Nital. Digital stereograph images were taken of each sample and fusion zone measurements taken using Pax-It quantitative image analysis program. Geometric dilution measurements have shown good correlation, within ± 2.6% (ref. Banovic 01) to ± 5% (ref. DuPont 99) with Electron Probe Micro Analysis measurements.Results and DiscussionTable 2 summarizes the geometrically determined fusion zone composition of each sample in conjunction with visual and microscopic observations.

In every instance copper was observed in the FZ, cracking was observed in the HAZ, most often along prior γ grain boundaries. The observations in Table 2 do not directly indicate a distinct composition range at which copper causes cracking the FZ of Steel – Cu welds, however, taking into account the margin of difference found by Banovic and DuPont, three regions of FZ crack susceptibility emerge: low, moderate and high. These regions are graphically shown in Figure 3.233Figure 6: Steel – 43 wt% Cu showing continuous copper films, some of which crackThere is very good agreement between cracking susceptibility and microstructure. Micrographs of varying compositions are shown in Figures 4 through 7. Figure 4 shows the FZ and HAZ of a Steel – 63 wt% Cu sample. This falls within the region of low FZ cracking susceptibility at high copper concentrations. No cracks were observed in the FZ, it is believed, because enough copper was present to backfill any cracks that might have formed.

No cracks were observed in the FZ, it is believed, because enough copper was present to backfill any cracks that might have formed. However, there was Cu penetration along former austenite grain boundaries. This is representative of copper’s effect on the HAZ when terminal copper was observed in the FZ. Figure 5 is a Steel – 49 wt% Cu sample, which falls within the region of high FZ crack susceptibility. Cracking in the FZ appears to fall along regions of continuous copper, which appear as copper “rivers”. As the sample cooled, thermal contraction stresses increased to a point that the copper film was no longer able to support the stress. This is classicalFigure 3: FZ Cracking susceptibility in Steel – Cu welds. green, Low; yellow, Moderate; red, High. Compositions of Figures 4-7 indicatedFigure 4: Steel – 63 wt% Cu. No cracking in FZ, but Cu penetration along prior γ grain boundariesFigure 5: Steel – 49 wt% Cu.

Compositions of Figures 4-7 indicatedFigure 4: Steel – 63 wt% Cu. No cracking in FZ, but Cu penetration along prior γ grain boundariesFigure 5: Steel – 49 wt% Cu. Crack in FZ extends into HAZ along prior γ grain boundary234solidification cracking. A Steel – 43 wt% Cu sample is shown in Figure 6. Cracking along continuous films of copper is easily seen. This falls well within the high crack susceptible region. Figure 7 is a Steel – 9.2 wt% Cu sample that falls at the lower composition cracking range, and within the moderate crack susceptible region. What is significant to note is the presence of copper appearing spheres on the order of one micron in diameter. These spheres were confirmed as copper by performing Energy Dispersive Spectroscopy using an EPMA in spot mode at 20 keV. The presence of these copper spheres at a composition less than the binary Fe- Cu peritectic isotherm maximum composition of Fe – 11.2 wt% Cu is not predicted by the equilibrium binary phase diagram.

The presence of these copper spheres at a composition less than the binary Fe- Cu peritectic isotherm maximum composition of Fe – 11.2 wt% Cu is not predicted by the equilibrium binary phase diagram. Their presence can be explained by any combination of three factors; 1) non-equilibrium solidification conditions, 2) ternary peritectic maximum copper composition that is less than 9.3 wt% copper and 3) composition determined by geometric dilution measurement was less than the actual composition. The simplest, and most probable given the high cooling rates associated with fusion welding, would be non-equilibrium solidification conditions.Figure 7: Steel – 9.2 wt% Cu. Small Cu spheres are seen in inter columnar regionSolidification ModelingIn order to determine the solidification condition, the dimensionless back diffusion parameter α was calculated following the Scheil (ref.) solidification model for the Steel – 7.2 wt% Cu sample.

Diffusion data was obtained for copper in δ ferrite (ref. Arita). Solidification time was estimated by using the Rosenthal (ref.) approximation to determine cooling rate and the binary Fe-Cu liquidus and solidus lines were used to determine the solidification temperature range (i.e. no liquid under cooling). Additionally, dendrite arm spacing measurements were obtained directly from the sample. The back diffusion parameter was calculated to be .00357, indicating that non-equilibrium, Scheil solidification conditions exist. It is believed that the presence of the copper spheres in the Steel – 9.2 wt% sample is a result of non-equilibrium solidification conditions.10.90.8d0.7i lo S n o i t c a r F0.60.50.40.3Knowing the solidification conditions, the quantity of terminal copper can be modeled using the Scheil equations in conjunction with a finite difference technique (ref.?).

To do this, the liquidus and solidus compositions were taken directly from the Fe-Cu phase diagram at each weight percent0.20.1013851335128512351085118511351435Temperature: deg C30 wt% Cu65 wt% Cu50 wt% Cu10 wt% CuFigure 8: Fraction solid vs. Temperature using Scheil model for various compositions of Copper235148510090copper from 1 to 100 wt%, resulting in a total of 200u C80l a n m r e T % a e r Ai70605040302010y = -8E-05x3 + 0.0214x2 - 0.2931x R2 = 0.9881measurements. The finite difference iterative technique using the Scheil equations was then used to predict the fraction solid vs. temperature for Steel – 10, 30, 50 and 65 wt% Copper as seen in Figure 8.00102030405060708090100Nominal Composition; wt% CuMeasured Area % CuScheil Model Area % CuEquilibrium Area % CuPoly.

(Measured Area % Cu)Figure 9: Area percent terminal copper versus nominal composition as predicted by Scheil and Equilibrium modeling, compared to measured valuesThe Scheil model was then used to predict the volume percent terminal peritectic copper phase and compare that to both the equilibrium predictions and the actual measured values of area percent terminal peritectic. Theequilibrium modeling used the Fe-Cu binary phase diagram for all data. The lever rule was used to determine the weight fraction of both the Fe rich and Cu rich phases at the Cu peritectic isotherm, 1096°C. Weight fraction was then converted to area percent by using knowing taking the average density of the Fe rich and Cu rich phases to be 7.96 and 8.96 g/cm3 respectively. It was assumed that area fraction and volume fraction would be equal, thus allowing a comparison between cross sectioned samples and solidification models.

It was assumed that area fraction and volume fraction would be equal, thus allowing a comparison between cross sectioned samples and solidification models. QIA was used to measure area fraction terminal copper in each sample by evaluating twenty fields per sample, resulting in the data displayed in Figure 9. High cracking susceptibility occurs when volume percent terminal copper is between 0.14% - 43%.In most situations, Scheil conditions will produce the greatest amount of terminal solute due to the absence of solid diffusion which does not allow the equilibrium solidus composition to be maintained throughout the solid with decreasing temperature. However, as seen in Figure 9, the equilibrium modeling predicts the greatest terminal solute. This is due to the decreasing copper concentration in austenite as Fe – Cu alloys cool. In the equilibrium case, diffusion of Cu in the solid would create a greater volume of terminal copper for any given nominal composition.

In the equilibrium case, diffusion of Cu in the solid would create a greater volume of terminal copper for any given nominal composition. However, in non-equilibrium Scheil conditions there is limited diffusion in the solid and the copper is retained in the solid matrix as the melt solidifies.The measured values show relatively good agreement with the Scheil model with the noted exception that at any nominal composition, less terminal peritectic copper was observed than the Scheil model predicted. This can be attributed to the effects of alloying elements other than Fe and Cu which constitute the binary phase diagram used for this modeling. Carbon content greater than .03 wt% has been shown to form two liquid layers in Fe – Cu alloys (ref. Iwase).

Carbon content greater than .03 wt% has been shown to form two liquid layers in Fe – Cu alloys (ref. Iwase). Furthermore, there are ??? elements in the Steel – Copper samples that may have a significant effect on the ability of the primary solid to absorb Copper.236SummaryA FGM of commercially pure copper into AISI H-13 tool steel was built using the LENS process. Solidification cracking and porosity were observed in the build. Dilution control experiments were carried out to determine the compositional cracking range of Copper in AISI 1013 steel, a representative system for H-13 tool steel. Cracking susceptibility was arranged into three categories; low, medium and high. Low susceptibility occurred from approximately 0- 8 wt% Cu and from – approximately 52.5 to 100 wt% Cu.

nctionally Graded Materials by Laser Metal Deposition 1Syamala R Pulugurtha, 1Joseph Newkirk, 2Frank Liou, and Hsin-Nan Chou3 1Department of Materials Science & Engineering, Missouri University of Science & Technology 2Department of Mechanical Engineering, Missouri University of Science & Technology 3Boeing Research & TechnologyAbstractFabrication of functionally graded materials (FGMs) by laser metal deposition (LMD) has the potential to offer solutions to key engineering problems over the traditional metal-working techniques. But the issues that need to be addressed while building FGMs are intermixing in the layers and cracking due to the residual stresses. This paper is to present the study of the effect of process parameters (laser power and travel speed) on the degree of dilution between the substrate (or, previous layer) and powder material for few metallurgical systems.

Keywords: LMD, FGMs, intermixing, cracking, dilutionIntroductionFGMs for industrial-related applications have been a subject of research for quite sometime now because of their ability to offer unique solution to the engineering problems over conventional materials and traditional composites. Among the various processing methods that are available, FGMs by LMD using solid free form fabrication route for obtaining bulk near net shape metallic components has been a promising technology. Its main advantage is the capability to control composition and properties within a fabricated structure by either pre-blending or combining different elemental powders using multiple powder feeder systems at the laser focal zone, which enables to tailor properties suited for specific engineering applications. This technology is ideal in producing functional prototypes and structural parts in smaller quantities and cost-effective [1,2,3].Functionally grading dissimilar/incompatible metals (i.e.

This technology is ideal in producing functional prototypes and structural parts in smaller quantities and cost-effective [1,2,3].Functionally grading dissimilar/incompatible metals (i.e. 100% A to 100% B) isqualitatively different from dissimilar/compatible systems [4] for the following reasons: (1) different thermo-physical properties, (2) composition becomes a parameter that can vary across the melt pool, and (3) nucleation of phases assumes importance, especially when inter-metallic compounds are present in the phase diagram. Although a lot of studies have been conducted on Ti-based systems by laser alloying, cladding, laser-rapid forming (LRF), etc. like Ti-N [5,6,7]; Ti-C-N [8], Ti-Al [9]; SiCp-Ti6Al4V [10]; TiC-Ti6Al4V and TiC+NiCrBSi –Ti6Al4V [11]; Ti- xV, Ti-xMo [12], a very limited literature is available on Ti based alloy/Ni FGMs [13], or Ti based alloy/ Fe FGMs.

The dissimilar systems such as Ni-Ti have following features: (a) different thermo-physical properties, for example, thermal diffusivity of Ni is roughly three times that of Ti at room temperature and density of liquid Ni is twice that of liquid Ti, and this significantly influences heat transfer. (b) There are three inter-metallic phases (IMPs) in the system which can form from the liquid through invariant reactions. Lin et al [13] successfully deposited a FGM having a continuous gradient from 100% Ti to 60% Rene88DT superalloy by LRF. They investigated the phase evolution during solidification and attributed the hardness of the graded material to be dependent on the amount and morphology of the phases: Ti2Ni, TiNi, and α + Ti2Ni eutectoid.

Most of the previous studies were focused on understanding the solidification behavior and phase evolution in multi-component systems, but the issues like intermixing in the layers, cracking due to thermal and residual stresses have not been dealt with. With this in mind, the present work focuses on the effect of process parameters like laser powerand travel speed on intermixing in the layers, and the pre-treatment (i.e. substrate heating) of the substrate to try counter cracking issues. ExperimentalOur LAMP system (Fig. 1(a)) consists of a 1 kW diode laser (Nuvonyx ISL-1000M, 808 nm, spot size 2.5 mm), a five-axis numerical control working table, and a powder feeder with co- axial nozzle was used to fabricate the graded materials. The experiments were conducted in an argon-gas environment to prevent the melt pool from oxidizing and oxide contamination from occurring during processing.

The experiments were conducted in an argon-gas environment to prevent the melt pool from oxidizing and oxide contamination from occurring during processing. The powder stream from the hopper was directly injected using argon gas into the molten pool through the laser nozzle. The metal powder was melted and subsequently re-solidified to form the clad layer (Fig. 1(b)). The typical size of the powders was in the range of 100-45 microns. The nominal composition of Fe-82 wt% V (powder-1) and Inconel-625 (powder-2) powders are listed in Table 1. The substrate materials used for the experiment were cold rolled 1 in. thick Ti6Al4V and SS316L.

The substrate materials used for the experiment were cold rolled 1 in. thick Ti6Al4V and SS316L. The substrates were cleaned prior to deposition using ethyl alcohol to remove any dirt or grease on the surface.Table 1: The chemical composition (wt%) of the powdersType of Powder Powder-1: Fe-82 wt% VPowder-2: Inconel-625Composition (wt%) V (82), Al (0.68), Si (0.9), C (0.07), S (0.01), P (0.02), Fe (18) Ni (58), Cr (20-23), Mo (8-10), Nb+Ta (3.15-4.15), Fe (5)(a)(b)Fig. 1: (a) LAMP system (fig. shows repair welding); (b) Schematic of the deposition process. (Source: http://www.emeraldinsight.com/fig/1560130405005.png)The as-deposited material was sectioned parallel to the compositional gradient for microscopy studies; for X-ray diffraction (XRD) analysis, the perpendicular section was utilized.

The composition along the gradient direction was characterized using scanning electron microscopy (SEM): Hitachi S570 equipped with a Si-drift energy dispersive X-ray (EDS) analysis facility. The identification of phases was achieved using X-Ray Diffraction (XRD)(Philips Xpert X-ray diffractometer), and micro-hardness of the deposits was measured using Struers (model: Duramin). Results and Discussion Composition and Phase AnalysisIn order to understand the effect of residual heat on inter-mixing/cracking with change of laser power on each deposited layer, a simpler experiment was designed by changing the power (Ramp up & Down) every 0.9th inch on a single layer track; parameters listed in Table 2. In case of powder-1, cracks were observed on both the substrates for all the processing parameters. The Fe-V equilibrium phase diagram (Fig. 2(a)) shows σ-phase is more likely to form in austenitic steels when there is ferrite retained from high temperature operation.

The Fe-V equilibrium phase diagram (Fig. 2(a)) shows σ-phase is more likely to form in austenitic steels when there is ferrite retained from high temperature operation. Also, V is a ferrite stabilizer and forms brittle σ-phase (needle like structure) under processing conditions. XRD phase analysis results further corroborates the presence of only BCC-Fe 2θ peaks in powder-1 deposit on SS316L, as shown in Fig. 3. The combination of brittle phases and residual stresses probably may have caused cracking in the powder-1 deposit on SS316L for all the processing parameters.Table 2: Processing parameters for different power profiles, * Ti6Al4V and SS316L substrates, powder-1 & 2Exp. TypePower, WTrvl. Spd., ipmPwd. Feed Rt., g/minRamp-UpRamp- Down300, 387.5, 475, 562.5, 650, 737.5, 825, 912.5, 1000 1000, 912.5, 825, 737.5,650, 562.5,475,387.5,300105(a)(b)(c)(d)Fig.

TypeReg. Type Substrate Dilution Deposit SubstrateSubstrateRegionTi Ni Fe 0.1 1.0193.41 54.63 16.36 92.37 59.2 42.62 27.25 93.47 58.59 40.83 20.271 2 3 1 2 3 4 1 2 3 4 1 2 3 1 2 332.5 60.23000.26 0.92 1.09 2.5 0.26 0.77 38.9 0.99 73.9 36.1 3.48 74.2 66.5 5.5429.6 40 48650DilutionTi64Deposit SubstrateRamp- up27 1.77 55.8 7.38 39.4 67.5 6.99 23.2 65.91000DilutionDeposit Substrate Dilution Deposit Substrate Dilution Deposit300SS316L1000Figure 2b shows the equilibrium phase diagram of Ni-Ti alloy. The three most important IMPs in the phase diagram are Ti2Ni, TiNi3, and TiNi, which can contribute to cracking, while Ti-Cr and Ti-Mo form solid solutions over the entire composition range with no IMPs. Attempts to deposit powder-2 on Ti6Al4V substrate resulted in cracks due to residual stresses at all the processing parameters, while no cracks were observed when deposited on SS316L, an example shown in Fig. 4.

4. A similar work carried out by Domack et al [14] showed macroscopic cracking in powder blends containing 40-60 percent Inconel 718 on Ti6Al4V substrate when produced by LMD. But Ni-based powders have been previously proven to bond well with steel in plasma- spraying, co-extrusion technique, or LMD [13,15] as there is no sigma/brittle phases present in the Fe-Ni equilibrium phase-diagram Fig. 2(c).Table 4: Parameters for laser deposition of powder-1 on Ti6Al4V substrateLaser Power, WTravel Speed, ipmPowder Feed Rate, g/min455535203677X-map results for powder-2 on Ti6Al4V and SS316L substrate show self-grading in themelt pool for both the set of experiments, as an example shown in Fig. 5. Elements like Al (- substrate) and Cr (-deposit) did not diffuse into one another, whereas Mo is uniformly distributed in the melt pool.

5. Elements like Al (- substrate) and Cr (-deposit) did not diffuse into one another, whereas Mo is uniformly distributed in the melt pool. The composition data in Table 3 shows that the laser power does not play a verysignificant role in controlling the inter-mixing in the melt pool. While that still holds true, laser power did result in segregation in the deposit. The deposit showed two discrete regions: (a) plume, and (b) non-plume. Plume- region consists of mixture of Ti+Ni, with higher weight percentage of Ti, while non-plume region has higher weight percentage of Ni. In general, Ni-Ti have relatively large negative enthalpy of mixing which results in the generation of additional heat in a very localized form in the melt pool, and hence aids the process of inter-mixing [16]. But segregation in the melt pool has shown that though inter-mixing was initiated, it remained incomplete during the deposition process at lower laser power levels.

But segregation in the melt pool has shown that though inter-mixing was initiated, it remained incomplete during the deposition process at lower laser power levels. At higher powers, the dilution zone is wide with more inter-mixing. Also, enthalpy of mixing results in formation of a much localized heat source in the melt pool, and can either raise or lower its temperature. If rate of solidification is considered to be the temperature difference between melt pool and the surrounding substrate material. A higher temperature of the melt pool from exothermic mixing will result in higher solidification rate. Figure 6 shows composition data from EDS analysis plotted on the Ni-Ti equilibrium phase diagram: the plume –zone with higher wt% Ti falls in the lower melting point regions (obviously will solidify the last) compared to the non-plume zones. Also, the density of liquid Ti is very low compared to that of liquid Ni (twice that of liq.

Also, the density of liquid Ti is very low compared to that of liquid Ni (twice that of liq. Ti) and therefore, the plume zone tries to move to the top along the solidification direction (Fig. 5). Because of the steeper slope of the liquidus NiTi line compared to that of Ti2Ni, the driving force for NiTi formation will exceed that of Ti2Ni. Therefore, NiTi will nucleate first (non- plume zone) from the liquid and Ti2Ni will most likely form heterogeneously on it through peritectic reaction (plume zone) [17]. Since the composition data lies close to the TiNi and Ti2Ni line compounds, it is safe enough to say that the possible cracking in the deposits is due to IMPs. In case of powder-2 on SS316L, we see only a small amount of Fe present in the deposit and no signs of cracks (Fig. 7 & Table 3).

Due to incomplete inter-mixing, the melt pool is divided into plume and non- plume regions.Fig. 6: Ni-Ti equilibrium phase diagram.(a)(b)CrCr120X60XFeNiMoFeNiMoFig. 7: Shows X-mapping images of RAMP UP power profile of powder-2 on SS316L at (a) 1000 W, and (b) 300 W. There is no inter-mixing in the melt pool.A multi-layered deposition of powder-1 on Ti6Al4V substrate was performed to study the effect of laser power on intermixing/cracking. Table 4 lists the process parameters. The X- maps in Fig. 8 at 455 and 677 W show some amount of un-melted powder-1 particles at surface of the deposit. The composition analysis (Table 5) shows self-grading in the melt pool, although the four layers are not distinct. But all the deposits showed cracks either due to residual stresses or formation of IMPs. Ti-V forms good solid solution with no brittle phases, but Fe-xTi has two stable IMPs (FeTi and Fe2Ti), and can contribute to brittle failure (Fig.

8: Shows X-images of powder-1 on Ti6Al4V substrate at (a) 455 W, (b) 535 W, (c) 677 W. The four-layers are not distinguishable and cracks penetrated through the entire deposit.(cid:0)Un-melted powder particle(cid:0)Macro-cracks(cid:0)InterfaceFig. 9: Macro-cracks in powder-2 on Ti6Al4V substrate, 600 W, 10 ipm, substrate heating @ 540oC.Inter-mixing and cracking in the layers for dissimilar materials are both inter-dependantproblems. In order to obtain a defect free structure (cracks, pores, or bonding defects) the substrates were pre-heated to 540 oC on a hot plate and maintained at that temperature during the entire deposition process. Table 6 shows the process parameters of powder-2 deposit on (with/without pre-heat) Ti6Al4V substrates. The results show that a pre-heat temperature of 540oC is not sufficient enough to eliminate the cracks formed in the deposits, an example shown in Fig. 9.

The results show that a pre-heat temperature of 540oC is not sufficient enough to eliminate the cracks formed in the deposits, an example shown in Fig. 9. Kelbassa et al [18] showed that a pre-heating temperature between 650-700 oC was required to obtain a defect free single LMD tracks for γ-TiAl deposit on Ti6Al4V and γ-TiAl substrates. A suitable pre-heating guaranteeing a defect free LMD result is still under investigation. Also, the X-maps (Fig. 12) show that at 300 W for with/without substrate pre- heating there is significant amount of un-melted powder particles observed at the surface, and this was also true at 600 W without pre-heating for all the parameters. Though exothermic enthalpy of mixing does contribute additional heat to the melt pool, the melting efficiency is strongly affected by processing parameters and material thermo-physical properties [19]. A dimensionless-parameter model was previously used to estimate the melting efficiency for the LENS deposits.

A dimensionless-parameter model was previously used to estimate the melting efficiency for the LENS deposits. Equation [1] defines melting efficiency as ratio of energy required for melting (AΔHm, per unit length) to actual absorbed energy (haP/S, per unit length).hm = Ch/Ry = SAΔHm/haP [1] Ch = S2A/α2 Ry = haPS/ α2ΔHmWhere S is the heat-source travel speed, A is the total deposit cross-section, ΔHm is the melting enthalpy, ha is the laser energy efficiency (which is assumed to be 40 pct [19], α is the thermal diffusivity of the substrate [20], and P is the laser power. For materials with dissimilar thermo-physical properties [powder-2 (2.67 J/mm3) onto Ti6Al4V (13 J/mm3) substrate], an average value of melting enthalpy between the two was used, 7.8 J/mm3 [19]. From the plot between Ch and Ry (Fig.

From the plot between Ch and Ry (Fig. 11(a)), melting efficiency can be determined solely from the dimensionless parameter Ry, for the material used in this study given by the following Equation [2]: hm = 0.1629 – (0.853/Ry) [2]At constant speed, the measured melting efficiency is slightly higher than that of the calculated value, Fig. 11(b). This discrepancy in the data can be attributed to the additional heat generated in the melt pool for dissimilar metals with exothermic enthalpy of mixing. The figure also shows how melting efficiency increases with the increase in the laser input power. It is previously reported that the calculated maximum melting efficiency for a moving point heat source in a 3D case is about 0.37 [20]. Also, Du pont et al [19] reported that increase in travel speed increases the melting efficiency as conduction can be neglected. In this case (Fig. 11(c)), the data contradicts the previously reported studies, which still need to be explained.

In this case (Fig. 11(c)), the data contradicts the previously reported studies, which still need to be explained. Pre-heating the substrates to 540oC before and during the deposition enhanced the inter-mixing in the melt pool for all the process parameters, as shown in Fig. 10. A significant amount of self-grading can be seen in the melt pool, as shown in Table 7. Geometric dilution ‘D’ which is measured between the melted substrate and deposited powder is given by D = As/(As+Ap) [3]Where As is the cross-sectional area of melted substrate and Ap is the cross-sectional area of deposited powder. The dilution increased with the increase in travel speed (Fig. 11(d)) as less power is delivered to the melt pool, resulting in decrease in deposit cross-sectional area. Also, dilution is affected by the melting efficiency. As the laser power is increased, more energy is available for melting the underlying substrate and incoming powder.

Also, dilution is affected by the melting efficiency. As the laser power is increased, more energy is available for melting the underlying substrate and incoming powder. Dilution was also observed to increase with substrate heating. MicrohardnessHardness was measured along the samples (Fig. 12(a,b)) and through the samples (Fig.12(c)). An average hardness value of 890±10 HV100 for powder-2 on Ti6Al4V substrate and 236±10 HV100 on SS316L substrate was obtained. A high hardness value in the former case may be due to the presence of brittle IMPs. The deposits were also tested at higher loads to study their performance. A small crack initiated in powder-2 on Ti6Al4V substrate at a load of 9.81 N (@ 5 s), while multiple cracks originated at 19.6N (@ 5 s). In case of powder-2 on SS316L, higher loads produced bigger indents indicating that the deposit is ductile/soft. Hardness as a function of depth for powder-1 on Ti6Al4V substrate at 1 N load, for 5 secs is shown in Fig. 12(c).

Hardness as a function of depth for powder-1 on Ti6Al4V substrate at 1 N load, for 5 secs is shown in Fig. 12(c). The profile can be divided into three distinct regions: clad, dilution and heat affected zones (HAZ). The hardness of the dilution zone initially shown to increase and after which dropped to value equal to HAZ. Such a sharp transition in hardness values can be potential reasons for cracks either due to residual stresses or IMPs.(a)300 W, 10 ipm NSH250 X(b)300 W, 10 ipm SH250 X(c)300 W, 20 ipm SH250 X(d)600 W, 10 ipm NSH250 X(e)600 W, 10 ipm SH250 X(f)1000 W, 30 ipm NSH250 XTiTiTiTiTiTiNiNiNiNiNihC(g)TiNi1000 W, 30 ipm SH250 XFig. 10: X-maps of powder-2 deposit on Ti6Al4V substrate under different process parameters: SH = substrate heating, NSH = no substrate heating.

An optimized window of processing parameters like laser power, travel speed and powder feed rate is yet to be determined to obtain a successful FGM. Inconel-625 deposits showed macro-cracks due to combined effect of IMPs and residual stresses on Ti6Al4V substrates, while no cracks were observed on SS316L substrates. The possible IMPs present in the deposits were NiTi and NiTi2. Intermixing and cracking was also observed on multi-layered (MLs) deposits of Fe-82wt%V on Ti6Al4V substrates. An attempt to counter the cracking issue in dissimilar materials by substrate treatment to 540oC did not solve the problem either. Substrate heating prior and during the deposition enhanced inter-mixing in the melt-pool. The geometric dilution (D) increased with increase in laser power, travel speedand substrate heating.

Substrate heating prior and during the deposition enhanced inter-mixing in the melt-pool. The geometric dilution (D) increased with increase in laser power, travel speedand substrate heating. Future work would involve optimization of parameters for these systems to reduce inter-mixing and cracking in the deposits, building MLs thin –walls and develop a simulation-model to study the melt-pool dynamics. AcknowledgementThis research was supported by a grant from the U.S. Air Force Research Laboratorycontract # FA8650-04-C-5704. Support from Missouri S&T Intelligent Systems Center is also greatly appreciated. References1. G. K. Lewis,U , E. Schlienger, Materials and Design, 21 (2000), p417. 2. S. M. Kelly, S. L. Kampe, Metallurgical and Materials Transaction A, 35A (2004), p1861.3. P.A. Kobryn, E.H. Moore, S.L. Semiatin, Scripta Materialia, 43 (2000) p299. 4. G.P. Dinda , A.K. Dasgupta , J. Mazumder, Materials Science and Engineering A 509 (2009), p98.5. M. Grenier, D.

nctionally Graded Polymer Matrix Nano-Composites by Solid Freeform Fabrication: A Preliminary ReportRichard Chartoff, Brian McMorrow, and Pierre Lucas Department of Material Science and Engineering University of Arizona; Arizona Materials Laboratory Tucson, Arizona 85712, USAABSTRACTA research program has been initiated to develop a Solid Freeform Fabrication (SFF) technology for combining nanosized particulate or fiber reinforcements with a photocurable thermoset matrix resin in order to produce functional graded composites. The composites that are being studied initially are optical components filled with nano-phase ceramic particles that form gradient refractive index lenses (GRIN). The Solid Freeform Fabrication (SFF) method employs an ink-jet deposition (IJD) process to form the composites.

The Solid Freeform Fabrication (SFF) method employs an ink-jet deposition (IJD) process to form the composites. The IJD process has the advantage of incorporating nano-reinforcements into a low viscosity matrix resin that is relatively easy to process and rapidly photocures to produce functional polymeric parts. It also has the advantage that major modifications to the basic SFF processing methodology are not necessary.The emphasis in the program is on demonstrating the feasibility of this approach for fabrication of gradient refractive index lenses (GRIN), which are flat instead of the traditional spherical lens geometry. As a result these lenses will be less costly to produce than conventional curved lenses. SFF is an ideal technique for meeting the needs of GRIN lens fabrication because changes in composition can be made from layer to layer and even within each layer, allowing for the introduction of compositional and structural gradients.

Thus it has the potential for creating the spatial material distributions required for designing computer optimized, custom made GRIN lenses. Integral to the SFF process are computer design procedures that specify the exact material deposition patterns that need to be employed in order to optimize the performance of the GRIN lens.The optical nano-composites will serve as a model system that we will use to work out the many challenges for implementing a viable SFF polymer composites technology. We then will make use of the information obtained and lessons learned from the work on optical composites and extend the development to structural composites that incorporate nano-particulate clays and carbon nanofibers.POLYMER NANO-COMPOSITES FOR GRIN OPTICAL APPLICATIONSBackgroundOptical lenses are traditionally created from blocks of uniform material by grinding curved (usually spherical) surfaces.

The spatial contour of the index of refraction change provides optical power when the shape and orientation are appropriately chosen. The non-planar surfaces (Figure 1) of this type of lens are detrimental to many commercial and military applications. The curvature is the source of the glint signature used for detecting covert observation. The curvature also will generate turbulence if the lens forms the window of a detector viewing out from a moving body, such as an aircraft. This situation often necessitates the use of an additional window fitted to the surface light curvature, which may adversely affect optical clarity or transmission.r) r ( nOptic AxisFigure 1 Traditional lens shape and refractive index profile.385Gradient index of refraction lenses (GRIN) present an alternative geometrical format. They provide the multi-dimensional spatial variation of refractive index required for optical power by implementing a non-uniform material distribution.

They provide the multi-dimensional spatial variation of refractive index required for optical power by implementing a non-uniform material distribution. (Figure 2) the use of GRIN techniques decouples the optical power generation from the curvature of the element. GRIN lenses can be fabricated with planar surfaces on either side or both sides. This should greatly simplify and reduce the cost of the fabrication of lenses by eliminating the need for grinding and polishing operations.The common methods usually used for generating a radial variation in the material properties are chemical diffusion or ion- bombardment, which implant foreign materials into a uniform cylinder or plate of a host material. The implantation density can be varied as a function of radial position. However, with these methods there are technical difficulties in achieving the required axial concentrations of material species.

However, with these methods there are technical difficulties in achieving the required axial concentrations of material species. Both of these methods also result in penetration- depth variations of the species introduced. This results in the optical properties changing as a function of distance along the optic axis. Thus independent control of the spatial variation of the material composition is not possible. As a result, the density profiles needed for optimal performance are difficult to achieve. So an SFF, layered manufacturing technique is an innovative, additive approach to the fabrication of GRIN lenses.r) r ( nOptic AxisFigure 2 Radial gradient GRIN lens shape and refractive index profile.Approach to Solving the ProblemIn our research program we are exploring the feasibility of an SFF ink-jet deposition process that creates GRIN lenses from polymer/ceramic composites that have composition tunable optical properties.

These composites are based on optically clear, photocurable thermoset resins containing various amounts of nanosized ceramic particles, to provide positionally controlled, tailored refractive indices. The utility of such composites in producing a wide refractive index tunability with little scattering has been demonstrated by Ho, et.al. [1]. The amount of light scattering in the composites is small, since the particles are much smaller than the wavelength of light.There are four major elements of our GRIN research. These can be delineated as:1. Polymer and composite development; this includes formulation or synthesis, ceramic particle dispersion, and polymer and composite processing Ink-jet printing and processing of layers;2. 3. CAD GRIN design implementation; 4.

3. CAD GRIN design implementation; 4. Process feasibility demonstration; characterization of optical propertiesCurrently we are concentrating on elements 1 and 2 (encompassing the more basic materials processing related issues) in order to validate the viability of the concepts proposed. We are using a standard piezo ink-jet printer head for the initial exploratory work to demonstrate the basic concepts outlined here. The reservoirs in the ink-jet printer head contain a range of polymer/ceramic concentrations, so that the concentration of particles (and the index of refraction) at any location can be controlled by depositing appropriate amounts of material from the different writing heads. Following deposition, the layers are photocured by UV light. The UV light application mode is controlled so as to provide suitable optical and mechanical properties in the final product.

Following deposition, the layers are photocured by UV light. The UV light application mode is controlled so as to provide suitable optical and mechanical properties in the final product. Aspects of polymer cure control will include optimizing the cure advancement of each layer precisely so as to achieve a seamless interface and the application of radiation in specific patterns that will minimize the effects of cure shrinkage and residual stress buildup in the cured material.386Ink-jet Printheads In order to achieve a sufficient refractive index gradient in a lens as large as several inches across, it will be necessary to both increase the material refractive index difference, Dn ( as discussed below) by using a higher index ceramic filler and to implement a more sophisticated print system. Present day advanced commercial ink-jet capabilities include high-performance, multichannel printheads.

Present day advanced commercial ink-jet capabilities include high-performance, multichannel printheads. For example, there are multi-nozzle piezo-driven printheads designed specifically for printing UV inks onto polymeric substrates at high deposition rates [ 2]. Nozzle diameters in these systems are of the order of 20mm or less. The printers also have gray-scale resolution capabilities with greater than 900dpi resolution. In order to work with higher viscosity formulations, the temperature can be controlled by a heating element integrated into the printhead mount that heats both the printhead and the ink supply. An example of the utility of printheads of this type in a related SFF application is micro devices made from a UV curable (cationic) epoxy resin [ 3].Printability of Nanoparticle Loaded Polymers; Nozzle Clogging The photocurable monomer that we have chosen to work with initially is hexanediol- diacrylate (HDODA). HDODA is a low viscosity liquid with viscosity of several cP.

HDODA is a low viscosity liquid with viscosity of several cP. At the low solids contents that will be used for this application ( << 20% by volume), the relative viscosity of the loaded polymer systems will remain low [4, 5]. This will insure ease of printability at low temperatures and help to mitigate the potential ink-jet printer nozzle clogging issue. That nozzle clogging should not be an issue is further suggested by the work of Sharp and Adrian [ 6], who considered particle bridging in particle loaded fluids flowing through microchannels. Their experiments identified 'shear-induced arching' as a mechanism that causes microtube blockage. They found that this mechanism is most likely to occur only when 0.33D < dp< 0.46D, where dp is the particle diameter and D is the microtube diameter. This was observed for flows of particle- laden fluids with concentrations as low as 0.5%.

This was observed for flows of particle- laden fluids with concentrations as low as 0.5%. Since in our case the ink jet nozzle is of the order of 20mm in diameter, dp is around 10-3 less than D.Table 1: Average Refractive Index of Some Common Ceramic Systems [ 7]Ceramic Filler Selection been Itthat appropriately large optical gradients can be achieved through the use of ceramic particles dispersed in a polymeric matrix [1]. Typical UV cured polymers have refractive indices in the 1.4 to 1.7 range.

Typical UV cured polymers have refractive indices in the 1.4 to 1.7 range. There are several ceramics that have a considerably higher refractive index than these organic polymer systems, as indicated in Table 1 [7].hasdemonstratedMaterialAverage Refractive Index 1.75 1.55 1.95 1.72 2.71 2.68 3.9 2.4 2.3 3.0 2.0Alumina (corundum) Quartz (SiO2) Zircon (ZrSiO4) Spinel (MgAl2O4) Rutile (TiO2) Silicon Carbide (SiC) Galena (PbS) Barium Titanate (BaTiO3) Zirconia (ZrO2) Iron Oxide (Fe2O3) Zinc OxideThe use of a high refractive index ceramic (>2.0) in an organic matrix provides for a wide range of refractive index in the GRIN optics. The refractive index in the graded optics can be controlled by varying the concentration (To avoid optical of the dispersed ceramic.

The refractive index in the graded optics can be controlled by varying the concentration (To avoid optical of the dispersed ceramic. scattering from the dispersed particles, the ceramic powders used in the GRIN system should be nano-particulate (5nm – 40nm) and should be very well dispersed.) A greater difference in the refractive index between the ceramic particle and the organic matrix will allow a wide gradient to be achieved with a minimum amount of ceramic particles. In calculating the387necessary GRIN optical parameters it is found that the larger is Dn, the easier it becomes to fabricate GRIN lenses with shorter focal lengths as well as thinner cross-sections.In summary, the polymer matrix will be of lower refractive index and the nano particulate ceramic will be of higher refractive index. The graded refractive index can be achieved by varying the ceramic loading in the matrix monomer as well as the deposition pattern and drop size for deposition.

The graded refractive index can be achieved by varying the ceramic loading in the matrix monomer as well as the deposition pattern and drop size for deposition. The initial matrix material chosen is HDODA, which, when polymerized, has a refractive index of ~1.6. Initial experiments will be performed using a ceramic such as zirconia due to its commercial availability and relatively high refractive index. After the feasibility of using nano-particulate zirconia has been demonstrated using the ink-jet printing method, the use of higher refractive index ceramics, for example, rutile, barium titanate, and iron oxide will be explored in order to increase the Dn window.We noted above that ceramic loadings of 20% by volume will be a maximum limit. Assuming a linear rule of mixing to estimate refractive index of ceramic dispersed systems (worst case), the effective refractive index of a 20% zirconia loaded HDODA system will be approximately 1.74.

Using appropriate optics design methods, the resultant Dn of 0.14 is found to be adequate for fabricating GRIN test optics of dimensions up to a few inches in size. For Fe2O3 the Dn is 0.28, which provides much greater latitude in GRIN fabrication. The initial goal for proof of concept is to fabricate an optical wedge with a linear radial refractive index gradient.Particle Stabilization and Dispersion A key factor in the printed optics fabrication of GRIN elements with ceramic particle index modifiers is complete deagglomeration and dispersion of the nano-ceramic particles. Agglomeration and poor dispersion of ceramic particles in the organic matrix will result in scattering, degrading the GRIN optic’s performance. Traditional deagglomeration and dispersion techniques are ultrasonication and surface treatment of the powder during synthesis [ 8]. Significant progress has been made in the synthesis of nano-particulate ceramic powders in recent years.

Significant progress has been made in the synthesis of nano-particulate ceramic powders in recent years. Spherical nano-particulate ceramic powders with narrow particle size distributions are commercially available. Significantly, the same sources that synthesize the nano-particle ceramic powders have been able to develop surface treatments to keep the powders deagglomerated and with a high degree of dispersability. Such surface treatments are applied to the powder particles during the synthesis process. The treatments are of various types, which effectively functionalize the surface with an adsorbed monolayer, resulting in steric-stabilization that prevents agglomeration of the powder. These coatings are tailored so that appropriate wetting of the particles by an organic monomer is achieved. The surface functionalized ceramic particles are dispersed in the monomer by mixing using an ultrasonic vibration technique.

The surface functionalized ceramic particles are dispersed in the monomer by mixing using an ultrasonic vibration technique. Because of the small size of the particles and relatively large surface tension forces, settling of the particles is not expected.Numerous references to steric stabilization of nanoparticles by surface functionalization by different approaches are available in the literature. Specific examples of this are contained in references [ 9-15]. These include grafting polymers such as silanes onto the particle surfaces [9-11, 15] or using adsorbed monolayers of a polymer or surfactant [13-15].Polymer Matrix: Photocuring and Residual Stresses; Chemical Stability; Air Bubbles A key issue in photocuring is that the effects of residual stresses must be minimal so as to distortion and spurious stress-induced refractive index effectively eliminate gradients.

This can be addressed in two ways: 1) by implementing random segmented cure patterns [16], where UV radiation is applied to small areas separated from each other, in a sequential pattern until the entire layer is cured, and 2) applying radiation in incremental doses small enough such that cure is carried out slowly. By combining these two principles thedimensional388polymer will retain rubbery characteristics long enough so that much of the cure-induced residual stress can relax before full vitrification of the resin occurs [ 17 ]. The use of incremental curing methods has been shown to be quite effective in minimizing residual stresses by allowing each small area to undergo a high degree of stress relaxation before being joined to other cured areas.The chemical stability of the proposed system is important because we wish to develop optical properties that are stable over long time periods. HDODA cures to produce transparent, water-white optical qualities.

HDODA cures to produce transparent, water-white optical qualities. The presence of residual photoinitiator quantities could cause yellowing to develop during aging that occurs over a period of time after the GRIN lens is In collaboration with a supplier of photoinitiators we have identified an effective produced. non-yellowing photoinitiator that is stable in the cured resin.In order to minimize dust contamination and reaction inhibition (due to presence of oxygen) of the samples, we are forming GRIN samples under a nitrogen blanket in a 'clean' environment. We do not anticipate problems with imperfections introduced through bubble formation. There are several reasons for this. First of all, previous SFF ink-jet deposition work [3] has shown that void free spherical micro -lenses can be formed with such printheads using a photocured cationic epoxy resin. Epoxy resins have higher viscosities than the HDODA formulations of interest here.

Epoxy resins have higher viscosities than the HDODA formulations of interest here. If an epoxy can be processed without bubble formation, it is most likely this will not be a problem with HDODA. Further we will be working with very thin layers (<25µm in thickness) and very small droplets. Also, there will be substantial overlap of droplets as the material is being deposited. All of these are factors that will promote dissipation of air bubbles. However, if we do find that micro-bubbles are a problem, we will add small amounts of an organic soluble surfactant to the resin formulations. This method of bubble suppression is very effective and is used frequently in chemical systems to prohibit bubble formation.Substrates The type of substrate to be employed for the proposed application will affect the face of the lens and might compromise the quality of light transmission. Three types of substrates will be considered as working surfaces.

Three types of substrates will be considered as working surfaces. In preliminary trial studies of the concepts proposed here, we formed thin films around .002" thick, using a conventional ink-jet printer with a mylar film as a substrate, (film similar to those used for overhead transparencies). The results indicated that both material deposition and the surface of the films produced were of good quality (by simple visual inspection). We have not yet looked at any quantitative measure of surface clarity, and roughness. Also, if mylar should not prove totally effective, we will work with silicon and sodium chloride single crystal substrates. Silicon has been used previously in RP fabrication of micro-ceramic components by other workers [18] and is commonly used for preparing AFM (Atomic Force Microscopy) specimens. We have used NaCl single crystals [19] as substrates for producing extremely smooth thermoset epoxy resin samples for AFM characterization.

We have used NaCl single crystals [19] as substrates for producing extremely smooth thermoset epoxy resin samples for AFM characterization. An advantage in using NaCl is that the substrate can be simply dissolved away from the deposited polymer after the polymer has been cured, leaving an ultra smooth surface behind.Previous Research on Inkjet Printing of Ceramic Filled PolymersInk-jet deposition methods have been used previously for SFF of ceramic particle-filled resins and have proved useful because of their improved resolution relative to other SFF methods. Relevant prior work of other research groups on SFF of ceramics using drop-on-demand ink-jet deposition is reviewed here to provide additional background information for those who would care to explore the topic in more detail. All of the previous studies concern SFF and deposition of monomers (and resins) containing high loadings of ceramic particles of appreciably larger dimensions than the nanoparticles of interest here.

All of the previous studies concern SFF and deposition of monomers (and resins) containing high loadings of ceramic particles of appreciably larger dimensions than the nanoparticles of interest here. The object in these former studies has been using the ceramic slurries to construct green-forms for making ceramic components. The resins in these compositions then serve as a binder for the ceramic powder, which is subsequently burned out prior to densification of the ceramic. Various references that concern ink-jet deposition of389monomers containing high- loadings of ceramic particles (greater than 40% by volume) are noted in the list of references [ 20-29].

Various references that concern ink-jet deposition of389monomers containing high- loadings of ceramic particles (greater than 40% by volume) are noted in the list of references [ 20-29]. These are representative of the work being done by three different research groups, those at the University of Michigan [20-22], University of Manchester (England) [ 21-23, 28,29], and UCLA [ 24-27].Additional research of note that relates specifically to GRIN lenses is that being done by the Cima-Sachs 3-D Printing group at MIT. A recent publication of theirs [30] discusses the use of the slurry based 3-D Printing process for fabricating complex structured materials by printing organic binders in selected positions on each printing layer. This process was modified in an effort to fabricate GRIN lenses by depositing polymers containing different concentrations of an alumina dopant at various positions in a silica powder bed.

This process was modified in an effort to fabricate GRIN lenses by depositing polymers containing different concentrations of an alumina dopant at various positions in a silica powder bed. Two different dopant concentration profiles, which had maximum alumina concentrations of 1.63 mol% and 2.50 mol%, were printed into the silica powder beds. The doped ceramic powders then were sintered at 1650oC into optical transparency, and the magnifying effects of GRIN lenses with gradient profiles of alumina were observed. This work is only in an exploratory state and the method requires extensive post- processing.

This work is only in an exploratory state and the method requires extensive post- processing. It will be interesting to learn whether sufficient resolution and control of concentration gradients for producing effective GRIN lenses can be achieved in this method, even though there are problems with spreading of the droplets due to the porous nature of the substrate [ 31].ACKNOWLEDGEMENTSPreliminary work on this topic was initiated at the University of Dayton's Rapid Prototype Development Laboratory in collaboration with Dr. Allan Lightman. We are pleased to acknowledge that the current project at the University of Arizona is sponsored by the National Science Foundation, Division of Design, Manufacture and Industrial Innovation under Grant No. DMI–0341924.REFERENCES1. P.K.H Ho, D. S. Thomas, R.H. Friend and N. Tessler, Science, 285, 233-236 (1999) 2. Refer to URL: http://www. xaar.co.uk for a complete description 3. W. Voit, K.V. Rao, and W.

nctionally Gradient Material Design and Modeling Using Hypertexture for SolidFreeform FabricationSeok-Min Park, Richard H. Crawford, and Joseph J. BeamanLaboratory for Freeform Fabrication The University of Texas at AustinAbstractSFF technologies have the potential to become manufacturing process that are capable of producing parts that have not been feasible by other techniques. The fabrication of Functionally Gradient Material (FGM) is one of the possible candidate. It is critical.to provide three dimensional material gradient data in the solid model to fabricate FGM. The approach is to model spatially varying material density distribution on implicit solid geometry using a hypertexturing scheme and a procedural volumetric modeling. It is essentially an extension of procedural solid texture synthesis, which has been effectively used to model cloud, gas, and flow stream. It will save the amount of information storage especially when the gradient pattern is repeating.

It will save the amount of information storage especially when the gradient pattern is repeating. Geometric operation becomes convenient since the material data are linked to the geometry only by the reference surfaces.1IntroductionSFF technologies have the potential to become manufacturing process that are capable of producing parts that have not been feasible by other techniques. The fabrication of Functionally Gradient Material (FGM) is one of the possible candidates. An FGM is a non-homogeneous solid which exhibits spatially varying volume fractions of the. constituents. It has its micro-elements systematically and continuously distributed and.controlled so as to provide functionality suited to the intended service environment. Layer based manufacturing becomes a natural choice when material distribution inside a part varies continuously in three dimensions. To accomplish the fabrication of FGMs, SFF processes must be supported by advanced software.

To accomplish the fabrication of FGMs, SFF processes must be supported by advanced software. The SFF community is currently experiencing a growing need for including additional information, such as material distributions in the solid model.For fabricating FGMs, it is critical to provide three dimensional material gradient data in the solid model. The material data should be able to exist in discrete, gradient, or mixed form. Predetermined spatial distribution data must easily be represented. in the data .format. The modeling tool should support procedural modification of the distribution interactively.2 Approach2.1 HypertexturingI used a hypertexture scheme to model spatially varying material density distributions. Hyper texture is a technique which is intermediate between shape and texture by using space-filling199applicative functions to modulate density.

Hyper texture is a technique which is intermediate between shape and texture by using space-filling199applicative functions to modulate density. The model is essentially an extension of procedural solid texture synthesis that is well known in computer graphics, but evaluated throughout a volumetric region using a Density Modulation Function (DMF) [4].: R3 -+ lO, ...,11 is defined for the solid geometry. All points x for which f(x) is zero are said to be outside the object. All points x for which l(x) is one are said to be strictly inside the object. Finally, all points for which °< l(x) < 1 are said to be in the object s fuzzy region. ThisIn hypertexture, a continuous characteristic function which is a mapping from fformulation gives the object surface an appreciable thickness.

ThisIn hypertexture, a continuous characteristic function which is a mapping from fformulation gives the object surface an appreciable thickness. Finally, the fuzzy shape and the solid become unified once solid texture functions are combined with the function that describes the object s fuzzy region [4].Procedural texturing or modeling methods can be categorized into two types: explicit and implicit methods. In implicit methods, a texture pattern is defined as a function 1 of points in the texture space, and the pattern consists of a level set of 1, that is, the set of all points at which the function has a particular value. Since implicit models tend to be continuous throughout a region of the modeling space, they are appropriate for continuous density and flow phenomena such as clouds and fog [4]. Thus, the implicit scheme will be chosen to model the material gradient because of their ease of specification and smoothly blending density distributions.

Thus, the implicit scheme will be chosen to model the material gradient because of their ease of specification and smoothly blending density distributions. The implicit density functions are best defined by summed, weighted, parameterized, primitive implicit surfaces [4].Implicit SurfacesAlgebraic implicit formulation was used as a representation scheme for both solid and sliced Implicit surfaces for solid geometry are two-dimensional, geometric shapes that geometries. exist in three dimensional space; they are defined according to a particular mathematical form. Intuitively, an implicit surface consists of those points in three dimensional space that satisfy some particular requirement. The requirement is represented mathematically by a function, generically named 1, whose argument is a point p. By definition, if 1(jt) = 0, then p is on the surface.

The requirement is represented mathematically by a function, generically named 1, whose argument is a point p. By definition, if 1(jt) = 0, then p is on the surface. The function 1 does not explicitly describe the surface, but implies its existence [1].Compared to parametric surfaces, implicit surfaces are receiving increased attention espe cially with· respect to their accurate yet compact depiction of solid objects and their innate blending properties. That is, implicit surface functions naturally describe the interior of an ob ject, whereas a parametric description of an object usually consists of piecewise surface patches and requires additional information for the interior. Geometric queries, such as point classi fication to determine whether a point is inside, outside, or on the surface, are simpler with implicit surfaces that with parametric surfaces.

Geometric queries, such as point classi fication to determine whether a point is inside, outside, or on the surface, are simpler with implicit surfaces that with parametric surfaces. The ability to enclose volume and to represent blends of volumes endows implicit surfaces with inherent advantages in geometric design and the corresponding fabrication.The union of two algebraic surfaces, unlike the union of two solid models, is usually given by the product of the corresponding algebraic functions. Intuitively, if 1 iszero, then any multiple of 1, including multiplication by another function, is zero. It is called the closure property. Unfortunately, the multiplication f112 confuses the sense of inside and outside. That is, points that are within both spheres as well as points that are beyond both spheres evaluate positively; only points within one and only one sphere evaluate negatively.Solid modeling emphasizes operations on volumes rather that on surfaces.

Accordingly, if a point isit typically defines the union of objects 11 and 12 by min(11' 12)'Intuitively,200within any sphere it evaluates negatively, regardless of the number of surrounding sphere. Conversely, max(fl' f2) evaluates positively if a point is outside any sphere, thus representing the intersection of the volumes.Analytic expressions approximating union and intersection are given as;uniona(fl, ... ,fn)intersecta(fl,'" ,fn)= (fla+,"', +r;;a) a (fl+"",+f~)~1where a > 0, lima --+ oo uniona(fl,'" ,fn) = min, and lima --+ oo intersecta(fl,'" ,fn) = max.2.3 Density Modulation Functions (DMF)The geometric gradient information is determined and controlled interactively by a DMF, which is used to modulate an object's density within its material space. It consists of the several procedural functions such as bias, gain, and noise functions which are the base level functions that higher order DMFs are built upon.

It consists of the several procedural functions such as bias, gain, and noise functions which are the base level functions that higher order DMFs are built upon. These functions are used to control some aspect of an object s spatial characteristics. The following briefly introduces these functions.The bias function, is mainly used to either push up or pull down an object's density around the middle of the fuzzy region. The bias function is typically defined by t In(O.5). The values at three points are fixed such that, bias(O) = 0, bias(0.5) = band bias(l) = 1. By decreasing or increasing b, the values in an object's fuzzy region can be biased up or down smoothly [12].In(b)The gain function can be effectively used as an intuitive method to control whether a function spends most of its time near its middle range, or, conversely, near its extremes. As a result, the density distribution can be tweaked to be either flatter or steeper across the fuzzy region.

As a result, the density distribution can be tweaked to be either flatter or steeper across the fuzzy region. The gain function over the unit interval, for example, can be defined as follows.gain(O)ogain(0.25)0.5 -9 2"gain(0.75)05+ Q 2 .gain(l)=1By controlling the value of g, the rate at which the midrange of an object's fuzzy region goes from 0 to 1,can be increased or decreased [12]. In case that the material information is supplied in a discrete data form, for example, directly from FEManalysis, a polynomiaLinterpolation. canbeexploited to. interconnectithe data points in a mathematical form forDMP. For gradedcom~ositions,analyticiJunctionsmustbe defined, capable of representingsmooth.variation over the domain of material subspace. There exists a number of effective polynomial bases for interpolation. One of the most popular bases is the Bernstein-Bezierbasis.

There exists a number of effective polynomial bases for interpolation. One of the most popular bases is the Bernstein-Bezierbasis. The basic formulations are as follows.pm nTensor: P(x, y, z) = L L L WijkBi(x)Bj(y)Bk(z) i=O j=O k=O Barycentric: P(x, y, z) = L Lm m-im-i-jWijkBi]k(X, y, z)i=O j=O k=O201B['jk(X, y, z)G~) xiyizk(l - x - y - z)m-i-i-kThe composition at a point can be considered as a blend of the control compositions with their influence determined by the value of their basis functions, analogous to the surface reconstruc tion with a meshof.control points.In order to represent objects made of rrruJtiple materials using hypertexture, the mathemat ical space must be modified. The space for the model must now include material subsets apart from the global spatial dimensions that captured the geometry and topology of the object.

The space for the model must now include material subsets apart from the global spatial dimensions that captured the geometry and topology of the object. The material subset also needs to provide its own space for geometry and topology because the hypersurfaces should be in primitive volume and it is convenient to separate homogeneous regions and heterogeneous regions.3 Results and Discussion3.1 Discrete Point Data SetI used a 64 equally spaced discrete point data set in three dimension for demonstration. This data set was scanned into the program to create an implicit polynomial form of continuous DMF by tensor product of trivariate Bernstein-Bezier interpolation. It was first mapped onto the pre-determined material subspace domain which contains the graded portion of the secondary material only. If discontinuous or isolated material distribution is needed, more material sub space would be created.

If discontinuous or isolated material distribution is needed, more material sub space would be created. One of the surface of the material subspace was mapped onto the surface of the geometry with the ·Figure 1 shows an example of material gradientdepth.in z on'eC1ClOJn.Figure 1: Trivariate Bernstein-Bezier interpolation3.2 Pattern SketchingWhen gradient information is based on iterative experiments, designers need to sketch and modify the gradient pattern in the solid model in accordance with the experimental results.I used the cascaded form of the Bias and Gain functions, which are controlled only by Bias Coefficient, b and Gain Factor, g. The mapping on the material subspace and the solid geometry followed the same procedure as the discrete data set representation.

The mapping on the material subspace and the solid geometry followed the same procedure as the discrete data set representation. The sketched material gradients with different Bias Coefficient and Gain Factor, are shown in Figure 2Figure 2: Material Gradient by Bias and Gain Functions3.3 ApplicationsThe first application chosen for demonstration, is the Willie Cube, a rounded cube with a sphere removed. It is defined by 1 - (a 2x 2 -+ a2y 2 -+ a2z 2 )-6 0.43 and b = 0.5. A sharply edged wedge occurs along each circular opening of the cuboid. As shown in Figure 3, this implicit surface was first polygonized and rendered using the polygnization scheme and the surface normal of each polygon.(b8x 8 -+ b8y 8 -+ b8z 8 )60, with aThe Bias and Gain functions were applied on the material subspace of the spherical coor dinates with given range of radius. The highest density at the inner surface of the spherical shell and zero density at the outer surface.

The highest density at the inner surface of the spherical shell and zero density at the outer surface. Figure 4 shows the material gradient represen tation and its sliced geometry that were mapped onto the center of the Wiffle Cube. Since both geometry and material data are formulated in implicit algebraic form, they retain the clo-203Figure 3: WifHe Cubesure property that T\rt~QD'r"TD the implicity of the representation even after geometric operations. Therefore, only a cutting plane, the slicing of the material and the geometry can be representedFigure 4: Spherical Material Gradient in WifHe CubeFGMs of ceramic and metal canbe··employed in advanced dental and medical applications, rnainlyas implants and replacements. As bioceramic materials such as Calcium Phosphate, favorably bioreact with existing bone, they can serve as porous media to support the ingrowth of new bone tissue,which results in permanent bonding with the body.

The strength and life of the implants can be even more enhanced by applying biocompatible metal such as titanium for the core of the implant.A simplified human femur implant for hip socket was created using blending and union of implicit spheres and elipsoids. The base shape was formulated in the following form. Detq,iledfeatures were added by feature based boolean operations. The geometry of the implant is rendered in Figure 5.Implant (x, y, z) blend=5.0 newx=0~5*x*1.414+1-0.5*y newy=0.5*x+O.5*(-2+y)*1.414tmp=pow(elipsoid(y,x,z +pow(elipsoid(newy, newx, z, 0.6,0.1,0.1), -blend) +pow(pow«y-3.4)*4/3,2)+pow«x-O.6J*4/3,2)+pow«z-0.0)*4/3,2), -blend)0.25,0.25),-blend)returnpow(tmp,-1/blend)-1.0elipsoid(x, y, z, a, b, c)result=x*x/a+y*y/b+z*z/c return (result <0.001? 0.001result)The same algebraic surface in implicit form was used for the material subspace.

The gradi ent region in the subspace is specified as a function of radius within the range of the material subspace. Therefore, The material distribution has constant depth of a primary material shell. Localized distribution also can be achieved by modifying the blob model for the •material sub space. The material gradient and the cross section along thez axis of the implant are shown in Figure 6.4 Conclusion and Future WorkI have successfully. implemented material gradient varying in. three dimension in implicit al gebraic surfaces. Material gradient region in the solid can be designed by generating implicit205Figure 6:J.VJ.CIJIJGJ"J.Q,J. vr'aOleIltfunction effective terpolated was chosen to information can be queried at any points without further interpolation.and J.J.J.(J~IJGJ. .1Q,.1J.J.J.\;'IJJ.J.VU is .very either in- .."... ...., .......

.1Q,.1J.J.J.\;'IJJ.J.VU is .very either in- .."... ...., ....... i- scheme materialWhile it is possible to model a general closed surface as a single implicit surface patches, higher-order algebraic surfaces become difficult to design because there is no readily perceived relation between polynomial coefficients and the shape of a surface. This has prompted the use of piecewise algebraic surfaces, also known as semi-algebraic sets or implicit patches. Each surface piece is low order and spans a particular cell, usually a tetrahedron, that is defined by a spatial partitioning. The formulation would become more efficient to generate complex shapes by employing implicit algebraic patches.References[1] Bloomenthal, Jules, et a1.

nctionally Gradient Material Representation by Volumetric Multi-Texturing for Solid FreeformFabricationSeok-Min Park, Richard H. Crawford, and Joseph J. BeamanLaboratory for Freeform Fabrication The University of Texas at AustinAbstractIn order to fabricate parts from Functionally Gradient Materials (FGM) by layered manufacturing methods, the SFF com- munity needs a method to represent material distributions in solid models. Gradient material distribution requires accurate and systematic representation, and must be compatible with existing geometric data. This paper presents a method, called Volu- metric Multi-Texturing, to represent a three dimensional density gradient by exploiting hypertexturing and volumetric density functions. This method utilizes procedural and implicit methods to design/acquire density information.

This method utilizes procedural and implicit methods to design/acquire density information. The implicit procedural approach, as opposed to an input database, allows users to interactively create and modify the design patterns without explicitly changing the values stored in the database. Further, it promises convenience in process planning, and efﬁciency in data storage and computation time. The theoretical approach, design procedure, and tool path generation for fabrication of an example part are presented in the paper. The design procedure presented is based on specifying the material gradient by surface modiﬁcation.1 IntroductionOne of the potential capabilities of Solid Freeform Fabrication (SFF) is production of parts that have not been feasible by other techniques. The fabrication of Functionally Gradient Materials (FGM) is imminently feasible in near future.

The fabrication of Functionally Gradient Materials (FGM) is imminently feasible in near future. Several researchers are endeavoring to fabricate FGM parts through creation of new methods or modiﬁcation of existing techniques [3, 5, 7, 13, 16].To support the fabrication of FGM parts, a new data exchange format must be created to transfer material in- formation to the manufacturing process. Most commercial solid and surface modelers represent parts as homoge- neous solids. Currently, these tools only provide support for attaching material information as a simple annotation to lumped regions for rendering and mass property calculations (e.g., mass and moment of inertia). Therefore, existing schemes are not sufﬁcient to deﬁne volumetric properties of parts.1.1 Related Workset), where a material dimension Kumar and Dutta [10] ﬁrst introduced the concept of a material subset, (r -sets is added to the spatial dimensions R through Boolean operations.

Recently proposed models for specifying are based on two approaches: interpolat- ing functions or discrete density data sets. Several trivariate function forms have been proposed to describe the spatially varying material gradient [11, 6]. These functions can be categorized as either parametric or implicit rep- resentations. Both types have advantages and disadvantages, depending on how they are to be utilized for material processing. As a three dimensional extension of surface discretization, voxel-based and volumetric mesh schemes to store material density data have been extensively researched. Pegna proposed a representation method of spa- tially varying data by extending volumetric meshes used for ﬁnite element analysis [12]. His suggestion was that the nodal point set in the FEA mesh to model the physical state of a part could be used to model a spatially varying material distribution as well.

His suggestion was that the nodal point set in the FEA mesh to model the physical state of a part could be used to model a spatially varying material distribution as well. This approach not only provides a means to store material data but also provides an interface to material gradient design using FEA [9, 8]. 350m3. The complete material space is constructed by combining these r2 ApproachThe major requirement of this work is providing access to exact values at any given spatial point inside the solid. To meet this requirement, we chose to explore an implicit procedural scheme. In this approach, a global implicit with material space deﬁned in a geometry space G the following conditions:3 is generally represented by a set of procedural functions Fan interior pointmaterial index number of materialsdensity at ~pdoes not explicitly or parametrically describe the pattern. However, the information does exist 3. Since the 3. Each ’s can vary throughout the part’s geometry.

However, the information does exist 3. Since the 3. Each ’s can vary throughout the part’s geometry. A single material region may require no sub-functions. Interior regions near corners or edges may be composed of many contributing sub-functions. These sub-functions interact with ’s are thus wrapped together as a single procedure, and each other as multipliers or by blending operations. The f the wrapped procedure is only visible to users in the form of level sets of F . The information ﬂow is graphically described in ﬁgure 1.The function Fiprocedurally and implicitly, and is obtainable by ~p material composition at any ~p must sum to 1, nquery of its density in the material space M functions Fmint’s are necessary to fully constrain M(cid:0)1intm’s that inﬂuence a query point ~pconsists of a number of sub-functions, f.

The number of required fFmmintmiiint() =d~pQuerying PointintpGeometry Space3GVolumetric Mappings e i t i t n E e c n e r e f e Rs e i t i t n E e c n e r e f e Re c a p Sf l a Hs e i t i t n E e c n e r e f e RMaterial Densitypd1 mFsubfunction 1fsubfunction 2f. . .subfunction nf. . .Material Space3MFigure 1: Information Flow of FGM Data2.1 Volumetric Texture MappingOur approach is based on volumetric rendering schemes by texturing in computer graphics [2]. Such schemes are very effective at representing fuzzy objects such as cloud and smoke. By analogy, FGM design is envisaged as the creation of material clouds in a conﬁned geometric space in structured and controllable manner. Another motivation for pursuing this approach is that, based on our research into expected applications, material gradients will be emphasized near the surfaces of a part.

Another motivation for pursuing this approach is that, based on our research into expected applications, material gradients will be emphasized near the surfaces of a part. This is because many FGM applications are meant to replace or improve the surface coatings on components designed for severe thermal and mechanical stress environments. Hence, from the user’s point of view, it is logical to optimize the design scheme for surface-to-interior material351gradients.The material spaces are related to the geometry only by reference entities. The reference entities can be any topological entities such as, vertices, edges, or faces. These topological members can be either structural or non- structural members. Non-structural members include non-manifold and multi-dimensional geometries as shown in ﬁgure 2. Most modern boundary representation (B-rep) kernels allow the coexistence of these members in their data structures.

Most modern boundary representation (B-rep) kernels allow the coexistence of these members in their data structures. In case a mapping needs to be deﬁned apart from the B-Rep structure, non-manifold members can be created anywhere inside the geometry and volumetric mappings applied to these entities.Non-Structural Topology MembersFigure 2: B-Rep Models with Non-Structural TopologiesThe material gradient in the part is procedurally designed and stored. The term ‘procedural’ was adopted to emphasize that the patterns are described by program code rather than by data structures. The major deﬁning characteristic of procedural data is that it is synthetic, generated from a program or model, not from a discrete data set and its interpolation.

The major deﬁning characteristic of procedural data is that it is synthetic, generated from a program or model, not from a discrete data set and its interpolation. Structured data sets can be included among the gradient representation in a procedural design by incorporating the gradient interpolated from the data as one of its primitive operations.2.2 Blending FunctionsVertex Pointnf3f1f2fFigure 3: A Part with Many Faces Joined at One VertexBlending functions are used to union multiple f ’s deﬁned on the surfaces of a part. Th number of the f ’s for blending can be from 2 to n depending on the location of joining. For example in ﬁgure 3, n faces are joined at one vertex point, while only two f ’s may be required at the joining edges for unioning the neighboring faces. At the352vertex point at which n faces meet, the corresponding f ’s need to be joined simultaneously.

At the352vertex point at which n faces meet, the corresponding f ’s need to be joined simultaneously. A blending function, based on the Ricci’s smoothing approximation of union operations [15], has been formulated to accommodate these situations. The function can be generalized into the following form:...............where a is a positive real number.The basic idea is to blend the overlapping portions linearly with respect to each other. The order of blend-ing is then controlled by a single parameter a. The smoothness is determined as. Figure 4 demonstrates the change of smoothness in blending of f1 and f2 as a varies.a = 1Surfacea = 21fS u r f a c e2fa = 10a = 80Figure 4: Blending of f1 and f2 with a2.3 AdvantagesThe procedural representation is very compact in storage because it needs to store only procedural instructions and their parameter values.

The size of a procedural gradient can be measured in kilobytes, while the memory size353of voxel-based gradient data may require on the order of megabytes. Efﬁciency in memory becomes particularly signiﬁcant when the material pattern is repetitive or uniform through the geometry, because the data for the mapping can be reused.Another advantage of the procedural method is that it has no ﬁxed resolution. In most cases, it can provide fully detailed information for any resolution required. Since a set of instructions procedurally deﬁnes the gradient information, the affected region can be edited simply by modifying the instructions or deﬁnition without visiting individual nodes or voxels. Dimensional changes of a part with procedurally described material information do not require redeﬁnition of the material space.

Dimensional changes of a part with procedurally described material information do not require redeﬁnition of the material space. Uniform or non-uniform scaling of the material space, therefore, becomes trivial.Compared to other representation schemes such as voxels, parametric formulations, and volumetric meshes, implicit procedural models appear to be more suitable for generating machine instructions for fabrication. For material delivery tool path generation, for example, material gradient samples are usually evaluated in an order determined by the process planning algorithm, not by the gradient design procedure. The implicit procedure ﬁts perfectly in such an environment because it is designed to answer a query about any point in the part at any time. Parametric or discrete models, however, lend themselves to path generation based on a ﬁxed sequence or discretization, which may not match the needs of the process planning algorithm.

Parametric or discrete models, however, lend themselves to path generation based on a ﬁxed sequence or discretization, which may not match the needs of the process planning algorithm. In most process process planning programs, using an explicit routine or discrete data requires running the gradient design procedure as a pre-process to generate the material gradien, which must be stored in a buffer for retrieval as necessary for rearrangement or interpolation during process planning. This reduces the efﬁciency of the process.3 Design ProcedureIn this section we demonstrate a possible design procedure to incorporate material gradient information in a given geometry. A FGM drill bit that we currently explorating for fabrication was chosen for demonstration. For this application, the customer wants more erosion resistance near the tip of the drill bit where wear is most likely to occur. Away from the tip toward the shank of the bit, fracture resistance becomes more important.

Away from the tip toward the shank of the bit, fracture resistance becomes more important. Tungsten carbide provides abrasion and erosion resistance, while cobalt maintains ductility and fracture resistance. The design procedure for this gradient pattern is divided into the following step: importing geometry, global gradient design, surface gradient design, and integration of all information.3.1 GeometryThe material gradient modeler is exercised on a boundary representation of the part. Commercially B-Reps are used almost exclusively in geometric modelers. Therefore, this scheme can be easily integrated with accessible formats such as ACIS [1], STEP, and IGES.Certainly, a B-Rep is not the only way of describing the shape of an object to a computer.

Therefore, this scheme can be easily integrated with accessible formats such as ACIS [1], STEP, and IGES.Certainly, a B-Rep is not the only way of describing the shape of an object to a computer. In academia, research continues into other methods, such as polyhedral models (e.g., the STL format popular in SFF), octrees, and implicit boundary models.Although there exist several other methods of deﬁning the shape of solid objects, a B-Rep provides several advantageous features to this project. Every geometric entity displayed on the screen is explicitly held in the model’s data structure, making interactive editing, interrogation, and annotation of objects straightforward. For example, the attachment of attributes to entities such as faces, edges, and vertices are very efﬁcient. Subdivision of geometric entities in a B-Rep is very meaningful in deﬁning the material space.

Subdivision of geometric entities in a B-Rep is very meaningful in deﬁning the material space. The regions that are emphasized for FGM design match well with B-rep entities.3.1.1 Importing Geometric Datathe ACIS 3D ToolkitTM, developed and commercialized by Spatial Technology Inc. was used extensively in this work. ACIS integrates wireframe, surface, and solid modeling by allowing alternative representations to coexist in an uniﬁed data structure. ACIS bodies can have any of these forms or any combination of them. Linear and quadratic geometries are represented analytically, and NURBS surfaces are used to represent free-form geometry [1]. Figure 5 shows the relationship between the topological elements that deﬁne the boundary representation of an ACIS model.354BODYLUMPSHELLSUBSHELLFACEWIRELOOPCOEDGEEDGEVERTEXFigure 5: ACIS Topology Tree for B-Rep [1]ACIS saves, or stores, model information in an ACIS save ﬁle called the ‘SAT’ format.

This format is widely accepted as a geometric data exchange format. Currently, most commercial solid modelers support the conversion from their own internal formats to the SAT format. Therefore, existing solid modelers, such as ProEngineer, SolidWorks, and CATIA TM, can be used for deﬁning part geometry, which can then be saved as a SAT ﬁle and imported in to ACIS for adding material gradient information.3.2 Design Global GradientA global gradient is used to impose an overall trend or pattern of material distribution in the part independent of any geometric entities. This information is stored at the BODY level of of the ACIS topology.

This information is stored at the BODY level of of the ACIS topology. If the gradient pattern and the geometry are relatively simple and the density distribution propagates in a single direction either linearly or radially, a global gradient in the material space is all that is needed to satisfy the design intent.Figure 6: Different Types of Global Gradients on GeometryGlobal gradient design basically exploits a three dimensional geometric mapping of implicitly designed material355space to the part geometry. The mapping is constrained by deﬁning skeletal points and the orientation of the global material space in the part geometry. Since the B-rep geometry already deﬁnes the half space of the part, the global gradient space is not necessarily required to conform to the shape of the part. Figure 6 gives simple examples of global gradient material space mapping to s geometry.

Figure 6 gives simple examples of global gradient material space mapping to s geometry. The ﬁrst example in ﬁgure 6 creates an ellipsoidal or spherical material gradient, while the second represents radial gradient in cylindrical coordinates. A simple one dimensional gradient can be designed by mapping an implicit cube that entirely covers the part, as shown in the third example in ﬁgure 6.3.2.1 Implicit Surfaces with Gradient TexturingAlgebraic implicit formulation is used as a representation scheme for the global material space. Implicit surfaces are two dimensional, geometric shapes that exist in three dimensional space; they are deﬁned according to a particular mathematical form. Intuitively, an implicit surface consists of those points in three dimensional space that satisfy some particular constraints. The constraints are represented mathematically by a function, generally denoted f , whose argument is a point ~p.

The constraints are represented mathematically by a function, generally denoted f , whose argument is a point ~p. In implicit methods, the gradient pattern consists of a level set of f , that is, a set of all points at which the function has a particular value. Since implicit models tend to be continuous throughout a region of the modeling space, they are appropriate for this application for their ease of speciﬁcation and their smoothly blending density distributions. The implicit density functions are best deﬁned by summed, weighted, primitive implicit surfaces.3.2.2 Gradient ProﬁleA gradient proﬁle that affects the entire gradient material space should be deﬁned in a one-dimensional parametric form, where the parameter ranges from 0 to 1. The curve can be formulated from spline curve ﬁtting of data points, or from a predeﬁned exponential sketching function, called the Density Modulation Function (DMF) (See [14]).

The curve can be formulated from spline curve ﬁtting of data points, or from a predeﬁned exponential sketching function, called the Density Modulation Function (DMF) (See [14]). The DMF consists of several procedural functions, such as bias, gain, and noise functions, that are the base level functions that higher order DMFs are built upon. The function is designed to roughly ﬁt a generic monotonic material gradient proﬁle. The shape of the proﬁle can be interactively modulated by more meaningful parameters than the coefﬁcients in polynomials. If the material gradient needs to designed by sketching a rough outline, a DMF is useful to initiate the design process.3.2.3 More Complex Global GradientsGlobal material space is not necessarily bound to a primitive shape. It can be shaped into more complex forms by performing Boolean operations.

It can be shaped into more complex forms by performing Boolean operations. Further, constructive implicit modeling schemes, such as skeletal models, can be imported to generate more complex material spaces independent of the geometry space. An example that uses a complex global material space is shown in Figure 7.3.3 Designing Surface GradientsOnce the global material space has been deﬁned, detailed surface gradients can be speciﬁed. This process may be considered as an extension of volumetric texture or hypertexture mapping used in computer graphics. The software will give a variety of choices in selecting the surfaces for material gradients. Users can choose surfaces from the set of sub-surfaces (FACE) of part, or the entire surface (SHELL) can be selected.

Users can choose surfaces from the set of sub-surfaces (FACE) of part, or the entire surface (SHELL) can be selected. Surface patches for the gradient can be created on the part if the topological entities do not provide appropriate surface area for grading.3.3.1 Gradient ProﬁleThe schemes used for the global gradient proﬁles can be applied to the design of proﬁles for surface gradients. Preferably, the proﬁle starts from the surface of the geometry and propagates into the interior of the part.3.3.2 Uniform Surface GradientIn a case where a surface gradient with the same proﬁle needs to be applied to all surfaces of the part, the proﬁle can be added to be the SHELL in the topological hierarchy.

This formulation becomes very concise and provides356Figure 7: Use of Implicit Skeletal Geometry for Global Gradientuniform control of the surface gradient.3.3.3 Face Speciﬁc GradientIf the surface gradient is applied on speciﬁc regions and matches subdivided FACEs in topology, the program supports this procedure in a very efﬁcient way. Each FACE entity can store surface proﬁle information. For example, the part shown in ﬁgure 8 consists of 11 FACEs. A gradient proﬁle can be designed on any combination of these FACEs. Each FACE may a contain different proﬁle or no surface gradient proﬁle at all.Figure 8: Selecting FACEs for Face Speciﬁc Gradient Design3.4 IntegrationOnce the global gradient and the surface gradients are deﬁned, interaction and blending properties need to be determined between the global gradient and the surface gradients, and among the surface gradients.

Unless the surfaces are coplanar and have identical proﬁles, the instructions for blending the material gradient from each surface must be deﬁned at the joining EDGEs or VERTEX.3573.4.1 Surface TrimmingThe surface gradient proﬁle acts as a multiplier of the global gradient to ensure the primary material density reaches zero at the part surface. The global gradient, which is not bound by the geometry when originally deﬁned, is now trimmed to the surface of the geometry by incorporating the surface gradient proﬁles.

The global gradient, which is not bound by the geometry when originally deﬁned, is now trimmed to the surface of the geometry by incorporating the surface gradient proﬁles. The overall material gradient is formed by the product of the global gradient proﬁle and the surface proﬁles as plotted in ﬁgure 9.10.9Surf V olumetr Ov0.8Volumetric Gradient0.7Surface TextureA'0.60.5SurfaceSurfaceAOverall Gradient0.40.30.20.1A0Figure 9: Surface Trimming by Surface Texture as a MultiplierThe ﬁnal gradient can be adjusted to the designer’s intent by modifying either the global gradient or each surfacegradient.3.4.2 Edge BlendingJust as chamfer, ﬁllet, and blend operations create new surfaces for joining two tangent discontinuous surfaces in solid modeling, a material chamfer or ﬁllet can be used to generate intermediate regions for blending the material properties of two different regions.

At every joining edge and vertex where more than two surfaces meet, the designer should specify the blending properties of the material spaces from each surface. The joining contour is simply shaped by changing the parameter a described previously. The value of a determines the tightness or the radius of curvature of the corner at the joining edges. A sharper corner is formed as a increases. The value of a can be as low as 1 and increased to inﬁnity for the desired sharpness.3.4.3 Vertex BlendingMultiple surface gradients can intersect at a vertex, while only two surfaces are allowed to meet at an edge. The blending of the gradients near a vertex also can be controlled by the parameter a. The same approach as the edge blending is used for this case.

The blending of the gradients near a vertex also can be controlled by the parameter a. The same approach as the edge blending is used for this case. A higher value of a will form a gradient shape more tightly bound to the reference geometry near the vertex.3.5 Design Change and ModiﬁcationThe procedure for designing material gradients is intended to be analogous to that of feature based modeling. In this way, the design of the material composition can be seen as an extension of geometric design. Therefore, modifying the design should be intuitive and user friendly.358A'Dimensional changes of the solid generally do not require redeﬁnition of the material space because the proce- durally designed material gradient adapts to the modiﬁed geometry based the mapping and proﬁling instructions.

When new entities are introduced to the geometry space by adding more features, new mapping, proﬁle, and blending information should be inserted in the material space.4 ResultsA FGM drill bit was designed using the Volumetric Multi-Texturing scheme. First, an ellipsoidal implicit material gradient was speciﬁed for the global gradient. Surface gradient proﬁles deﬁned by DMFs were then added to all 11 faces. Material ﬁllets or chamfers were applied at all the joining edges and vertices. The ATTRIB class, which is derived from the ENTITY class on every ACIS topological entity, was exploited for storing material design information. Therefore, no auxiliary data structures were necessary to store material gradient data. Figure 10 shows a cross section of the FGM drill bit. The tungsten carbide rich region is located near the tip and the shell of the part.

Figure 10 shows a cross section of the FGM drill bit. The tungsten carbide rich region is located near the tip and the shell of the part. The composition gradually changes from the tip to the shank.Figure 10: A FGM Drill Bit, (Blue: Tungsten Carbide, Red: Cobalt)Further, tool path generation for multiple materials was implemented to demonstrate the feasibility of the scheme for SFF manufacturing. Closed loop contours of given material ratios were iteratively generated until they ﬁll up the entire cross sectional area of the drill bit. The path planning algorithm used the marching square algorithm to ﬁnd material isocontours on a 1000 by 1000 grid. We assumed that the SFF process would use 10 different material compositions from 0% to 100% of the primary material density. The result is shown in ﬁgure 11.

We assumed that the SFF process would use 10 different material compositions from 0% to 100% of the primary material density. The result is shown in ﬁgure 11. It was successfully produced without running pre-processing for rearrangement or creating a buffering datastorage.5 Conclusion and Future WorkAs a part of our research to produce FGM components by Selective Laser Sintering, a theoretical approach and its implementation for material design was introduced in this paper. A formulation responding to the initial spec- ; was achieved using procedural implicit schemes in conjunction with the iﬁcation, F Volumetric Multi-Texturing.3ii2(~pint) =d;~pintGmpWe have designed a volumetric material gradient inside a drill bit as a demonstration. The material gradient was combined with geometry created by a commercial solid modeler.

The material gradient was combined with geometry created by a commercial solid modeler. The gradient was composed of one global359Figure 11: Multiple Loop Iso-Contour Tool Path for FGM Drill Bitgradient and 11 surface gradients with joining and trimming operations. Both creation and modiﬁcation of the ma- terial gradient were very systematic and concise. Tool path generation from the formulation showed its feasibility for SFF manufacturing.We believe that our approach is a very efﬁcient tool for designing material gradients in a part from a given geometry. To ensure the robustness of the approach, we suggest several areas of further research. Designers may have tools to numerically simulate and optimize the material distribution for analytical material gradient data. Many researchers are exploring this aspect of FGM design [8, 9].

Designers may have tools to numerically simulate and optimize the material distribution for analytical material gradient data. Many researchers are exploring this aspect of FGM design [8, 9]. However, even when the material data is numerically optimized for the intended purpose, the designer may wish to adjust the results for actual fabrication. Our approach will become a more attractive tool if discretized volumetric data can be converted to the Volumetric Multi-Texturing representation for convenience in modiﬁcation and process.AcknowledgmentWe gratefully acknowledge the ﬁnancial support of the National Science Foundation under grant number DMI- 9618034.References[1] Corney, J., “3D Modeling with the ACIS Kernel and Toolkit”, John Wiley & Sons, West Sussex, England,1997.[2] Ebert, D. S., et al. (eds), Texturing and Modeling: A Procedural Approach, AP Professional, Chestnut Hill,MA, 1998.[3] Fessler, J. R., Nickel, A. H., Link, G., Prinz, F.

nctionally Optimized Ceramic StructuresC. Gasdaska, R. Clancy, M. Ortiz# and V. lamalabad AlliedSignal, Inc. Research and Technology Morristown, Nl 07962Anil VirkarDepartment of Materials Science and Engineering, University of Utah,Dragan Popovitch Advanced Ceramics Research Tucson, AZThe feasibility of using the Fused Deposition of Ceramics (FDC) process to rapidly fabricate functional quality advanced ceramic components has been demonstrated Multiple extrusion heads enable the deposition ofspatially engineered ceramic microstructures on the scale of250 pm.. This unique capability ofFDC allows components to be built with combinations ofmaterials and properties that are difficult or impossible to produce using conventional fabrication processes. Some concepts will be presented, along with examples ofmultiple material laminates produced using FDC.

Some concepts will be presented, along with examples ofmultiple material laminates produced using FDC. Strength data will be presented which demonstrates the performance improvement possible using spatially engineered microstructures.IntroductionBy suitably combining materials, it is possible to achieve improvements in properties such that the strength and toughness of the combination is greater than that for either material alone.

Two multi-material approaches used for strengthening and flaw tolerance in brittle materials are development of residual surface compressive stresses due to coefficient of thermal expansion (CTE) mismatchl composite structures2, and fabrication of silicon nitride/ boron nitride fibrous monolith.The concept of using combinations of ceramic materials with different coefficients ofthermal expansion (CTE) to produce residual stresses on cooling from the fabrication temperature is well documented in the literature l produce residual stresses have also been studied3 observed present work, we are illterested in pursuing multi-material combinations which are com.patible and co-sinterable with silicon nitride, the leading candidate for high temperature structural applications such as gas turbine engine components. The I-dimensional aspect of part building using FDC allows for.structures more complicated than the simple laminates prepared.using traditional methods4.

The I-dimensional aspect of part building using FDC allows for.structures more complicated than the simple laminates prepared.using traditional methods4. Systems which rely on phase changes to . Strength increases of 300% have beensome of these systems. Reliability and flaw tolerance is also improved. In theFibrous monolith (FM) materials are essentially unidirectional composites in which thevolume fraction ofthe matrix material is < 20%. The most common example consists ofa# Located at AlliedSignal Engines Division, Phoenix, AZ705Silicon nitride polycrystalline cellCell boundaryFigure 1.

The most common example consists ofa# Located at AlliedSignal Engines Division, Phoenix, AZ705Silicon nitride polycrystalline cellCell boundaryFigure 1. Cross-section ofan idealizedfibrous monolith microstructure.structure containing silicon nitride filaments surrounded by a sheath of boron nitride (see Figure 1).Cracking in these materials tends to be confined to the boron nitride layer, resulting inextensive delamination under bending stresses2 investigated which would allow a greater amount of fiber pullout and work offtacture under tensile loading conditions. Typically, FM materials are produced using filament-winding techniques using fibers which have the FM structure. The FDC process offers a more cost effective method for producing these materials..

Typically, FM materials are produced using filament-winding techniques using fibers which have the FM structure. The FDC process offers a more cost effective method for producing these materials.. Alternatives to boron nitride are beingThe solid freeform fabrication process we are using is based on the Stratasys machine.This technique uses solids loaded filament to deposit roadsofa binder/ceramic powder mixture4 • For multi-material parts, we are using a Stratasys 1650 dual-liquefier rapid prototyping machine.LaminatesIn order to take advantage of the capability of FDC to spatially distribute materials, systems must be found which not only result in desirable material property combinations, such as residual stress production, but which are also capable of being co-sintered without excessive cracking or delamination. Toward this end, we have been examining the use of particulates dispersed in a silicon nitride matrix to increase the CTE.

Toward this end, we have been examining the use of particulates dispersed in a silicon nitride matrix to increase the CTE. Two candidate material systems for use in silicon nitride are listed in Table 1. Material Melting/Decomposition Coefficient ofYoung's Modulus (GPa)Density (g/cmJ)Tempertaure (OC)thermal expanSIOn (fC x 106 ) 4.3 9.35SiC TiN Table 1. Propertles ofcandldate partlculate dlspersants.2700 2940448 2603.22 5.22706In order to evaluate the effectiveness of these materials in strengthening silicon nitride,bi-materiallaminates were prepared using slip-casting. The materials consisted of silicon nitride and silicon nitride plus up to 30 weight% of either TiN or SiC. The samples were sintered and dense samples were successfully prepared. Test bars were machined and room temperature bending strength measured. Results for samples containing SiC appear in Figure 2. The geometry of the samples is also illustrated.

Test bars were machined and room temperature bending strength measured. Results for samples containing SiC appear in Figure 2. The geometry of the samples is also illustrated. The resultsindicate that average strength can be increased by up to ~27%. In addition~ the decrease in the arnountofstrengthening at 800 °c is consistent with the drop in residual stress as the test temperature approaches the temperature at which stresses start to set on cooling.Verification that the increase in strength was due to residual stress was accomplished by amethod of progressive material removal and strain measurement using strain gauges. In this technique3 a strain gauge is mounted on a laminated sample and the build-up of strain is monitored as material is progressively removed ftom the opposite surface. A residual stress of ~157 MPa was measured using this technique. The differential strain between laminate materials180170160 150'(jj ::::=.- ..c:.... Cl c: (J)...

A residual stress of ~157 MPa was measured using this technique. The differential strain between laminate materials180170160 150'(jj ::::=.- ..c:.... Cl c: (J)... 1il -c c: (J) .c140130-c:'0 a.120I C')SN-20/80-SN............."..... ~SN-10/90-SN laminate.1~rnSEIte"·°1o 10/90oSN-30/70-SN laminate............ ""~. .f · 30170 ".SN-30170-SN laminate (800°C)'V'VSN'V..... '1.~ · .. ········0·····'V110o100SN= Si3N4 + sintering aids (see Table 1) 10/90 =90wt% SN+ 10wt% SiC (composition 1) 20/80 = 80 wt% SN + 20wt% SiC (comp. 2) 30170 = 70wt% SN + 30 wt% SiC (comp. 3)k::::-23 mm>1Geometry of 3-point specimen used for testing laminates.Figure 2 Strength data obtainedin 3-pointbendingfor laminates prepared from the indicated compositions. Datafor boththe laminates and the "core" materials are presented and connected by the lines.

Datafor boththe laminates and the "core" materials are presented and connected by the lines. Individual data points are plotted, with the average strength for each group represented by the filled diamond symbol. In all cases" the average strength ofthe core (10/90, 20/80 or 30/70) is less than. thirlaminate· strength,if1dica;t~ngistr¢ngthenirtgduetosurface compressive stresses. The drop in strength at 800c>C is also consistent with the presence ofresidual stress due to thermal expansion mismatchbetween the core and outer skins. The strength ofthe outer skin by itselfis shown by the group labeledSN707was measured as 5.4 x 10-4 permanent stresses begin to set in at 1300 °c, then a difference in CTE of 0.42 x 10-61°C would produce a differential strain of ~ 5.4 x 10-4Ifwe assume that on cooling from the sintering temperature.Using combinations of materials· with CTE mismatch to produce residual stresses raises anumber of issues.

For example, residual stresses must not be so large as to induce<spalling or delamination between materials. In addition, the bonding at the interface and the presence of interface defects are important characteristics in determining the ultimate degree of strengthening. The spalling and delamination question is strongly dependent on the interfacial bonding and Mode II fracture energy at the interface. Techniques for measuring interfacial fracture energy are under development. However, for the problem of failure from the surface versus failure at the interface - where residual tensile stresses develop - has been outlined for the case of bending. As an example, assuming a three layer laminate stressed in bending the corresponding fracture strengths (in flexure) are:dzE~Eo o O'P(s) = O'sP + (1 - v)dfor failure from the surface, andO'P(I) =d ° d2dlE~Eo O'IP - (1 - v)dzzfor failure from the interface.

Where, O'~F and O'YF are the true fracture strengths for failure from surface and interface flaws, respectively, dl is the outer layer thickness, d2 is the inner layer thickness, d = 2d l + d2 and ~E is the free strain difference between materials (related to the difference in CTE and ~T). E is Young's modulus.The preceding equations show that to maximize the strength as well as the damage resistance, the following are required: (1) high O'°IF: This can be achieved by improving processing such that interior flaws are kept to a minimum. (Failure from the interior may not be bad either); (2) the inner layer should be fully encapsulated in the outer layer such that there is no chance of introducing damage in the inner layer during service (complete confinement); (3) enhance ~Eo to as high a value as possible such that a high surface compressive stress exists. This provides damage resistance.

This provides damage resistance. However, the surface compressive stress should not be so large as to cause delamination; (4) decrease d 1 to a value small enough (relative to d2) such that interior tensile stress is kept toa minimum, yet dl is large enough to prevent contact-induced damage from penetrating too deep into the.inner region. Preliminary calculations show an outer layer of thickness ~150 to 250 Jlm to be more than adequate for most applications. For a component of 5 mm thickness, the d2/2dl ratio is then between 9 and 15.67. For a component of 20 mm in thickness, with an outer layer thickness of ~500 Jlm, the ratio is 19. The corresponding tensile stress in the inner layer is very small.

For a component of 20 mm in thickness, with an outer layer thickness of ~500 Jlm, the ratio is 19. The corresponding tensile stress in the inner layer is very small. This latter example corresponds to the case of the SiC insert being examined by AlliedSignal Engines (see next section), where the insert is located ~ 380 microns below the silicon nitride surface and the overall thickness is ~ 20 mm.708Design using FDCWhile the preceding work has demonstrated the feasibility of using compatible, co-sinterablematerial combinations for strengthening, the ultimate goal is to use the unique spatial control capabilities of FDC to design and build components. As part of the project on solid freeform fabrication, the Engines division at AlliedSignal (ASE) has been working on designing a multi material turbine blade which uses residual stress strengthening.The design used for modeling consists of a ceramic blade inserted in an AF-2 alloy disk (Figure 3).

In this initial work, design analysis was focussed onthe attachment area since this region is known to experience the highest stresses and contains the area most likely to fail. Stresses in the attachment area were analyzed for a rotational speed of 33,500 RPM and a uniform temperature of 1450 OF . Based on these conditions a uniform stress of24.1 ksi was imposed outside the attachment area. The coefficient of friction between blade and disk was assumed to be 0.6. This was based on previous experience with ceramic blade attachment systems. Using these inputs, a baseline condition was analyzed. The peak stress was calculated to be 105 ksi and located just outside the contact area between the blade and disk. It arises from the high frictional load between blade and disk.Various design strategies were evaluated to reduce this peak stress and are illustrated in Figure 3.

It arises from the high frictional load between blade and disk.Various design strategies were evaluated to reduce this peak stress and are illustrated in Figure 3. Although a silicon carbide material has been modeled as the higher CTE component, particulate based silicon nitride composites would provide nearly equivalent behavior. Three of the four geometries examined utilized combinations of SiC and SbN4 to produce residual compressive stresses in the contact area. The geometry illustrated in Figure 3d - a blade containing a central cavity to reduce weight - actually caused a 77% increase in peak stress. Figure 3c illustrates the most promising approach. A silicon carbide insert in the contact area significantly reduces the peak stress.t -~<;+--AP-2as"SiC/ /\';'1 '! /1 /01,[~"jb.015AF-2clJdI IFigure 3.

----i1--.-- - - - - -- - - - - -0.07009001100 1300 Set temp., deg C15001700Figure 4. Variation in the maximum tensile stress in the blade attachment area as a function ofset temperature.Fabrication of a complete blade using multi-material FDC is expected to begin shortly.In the meantime, the feasibility of the process has been demonstrated in sub-element testing. Figure 5 shows a cross-section of a turbine blade with a narrow band of insert material at the attachment area.Figure 5. Cross-section ofb a e built with insert material in the attachment area (black line at right).Fibrous MonolithsA major concern regarding the fused deposition of filaments with fibrous monolith structure was whether or not the fibrous monolith structure would be retained as the filament was melted in the liquefier and extruded.

In order to address these concerns, Advanced Ceramics Research (ACR) fabricated a fibrous monolith filament containing a single core and sheath, using a model system containing alumina and carbon black for this evaluation. The core material was RU9710(Stratasys, Inc.) binder containing 55 vol% alumina while the sheath consisted ofRU9 ~inder containing 42.75 vol% alumina and 2.25 vol% carbon ?lack.. Th~c?re and sheath constituted 82.5 and 17.5vol% oftb.~ 0.070 inch diameterfilament,.r~$pectiyely.A number of test pieces were produced and a crpss-sectionof one piece is shown in Figure 6. In ordertotnor¢clearly discern the stmcture,the roads inthispiec~were laid down on0.44 mm42.75v/o A1203, 2.25 v/o C-blackFigure 6. Cross-section ofsingle-cellfibrous monolith sample prepared byfused deposition offibrous monolith filament.top of e(lchother, resulting in the voids which appear in Figure 6.

Cross-section ofsingle-cellfibrous monolith sample prepared byfused deposition offibrous monolith filament.top of e(lchother, resulting in the voids which appear in Figure 6. As seen in Figure 6, the fibrous monolith microstructure Was retained during the deposition process. A 0.015 inch diameter nozzle was used for deposition, resulting in a road width of ~ 0.017 inches. ACR also prepared a 0.070 inch diameter filament containing 12 cells, each with a core and sheath, of alumina and carbon black respectively.A micrograph of the. filament cross-section iS$hown in Figure 7. The compositi?n of the cell and cell walls was again alumina and aluminalC-black, respectively. The composition of the cell boundary material was adjusted slightly: the solidsloading was increased from 45 to 54 volume % solids. Stearic acid waS used as the dispersantfor both cells and cell walls.The cell structure retention is excellent after extrusion.

Stearic acid waS used as the dispersantfor both cells and cell walls.The cell structure retention is excellent after extrusion. Also shown in Figure? is across-section of a sample built with some overlap between adjacent rows... Thisoverlapis commonly used in FDCto avoid the introduction ofstrength limiting voids. Thoughsome loss of structure occurs when roads overlap during a build, the overall retentioll of the FMstructure is quite good. Deposition parameters will need to be controlled closely to avoid excessive. smearing under these conditions. However, it appears that the microstructural scale o],tained (5 micron cell wall thickness) using single cell filament will be sufficient to get theadvatltages of the fibrous monolith microstructure.711. - - 1.9mm-+...1.78 rom ----..Figure 7.

- - 1.9mm-+...1.78 rom ----..Figure 7. Micrographs ofmulti--cellfibrous monolith model material showing the structure inthe as-preparedfilament (left), after extrusion through a 0.3 mm nozzle (center) and after deposition with a O.076mm overlap between roads (right).The work on FM structures is being extended to materials based on silicon nitride as the core material and lanthanide glasses· for the cell walls. Some promising lanthanide·glass compositions have been prepared and will be used to manufacture filament in the next year.AcknowledgementsThe support of DARPA under contract NOOO 14-97-C-0115 is gratefully acknowledged.Also, we would like to thank Sriram Rangarajan of Rutgers University for fabricating the part illustrated in Figure 6, and Don Baskin for his assistance in the preparation ofthe laminates.References1. R. Sathyamoorthy, A. V. Virkar, and R. A. Cutler,J. Am. Ceram. Soc., 75 [5] 1136-1141 (1992). 2. D. Kovar, B. King, R. Trice and J. Halloran, J.

ndamentals of Liquid Phase Sintering During Selective Laser SinteringD. E. Bunnell*, S. Das*, D. L. Bourell*, J. B. Beaman*, and H. L. Marcus"tCenter for Material Science & Engineering, University of Texas at Austin, Austin TX 78712 tInstitute for Materials Science, University of Connecticut, Storrs CT 06269-3136IntroductiQnOne of the advantages of the Selective Laser Sintering (SLS) process is that a variety of materials can be processed. However, the goal of being able to produce fully dense metal parts with no post processing has been elusive. Using Selective Laser Sintering to produce metal parts with full density without post processing poses a challenge since both the processing conditions and the metal system must be controlled. This article describes two metallurgical mechanisms by which loose metal powder beds could be sintered to nearly full density using a scanning laser beam.

This article describes two metallurgical mechanisms by which loose metal powder beds could be sintered to nearly full density using a scanning laser beam. The mechanisms are particle rearrangement during liquid phase sintering (LPS) and in-situ infiltration. Some of the particles, when heated by the laser radiation, melt and form a liquid. If this liquid has certain physical properties (e.g., low viscosity and high surface tension) and wets the other solid particles, then the SLS process can in theory produce dense layers by either mechanism. The purpose of this study is to determine the process and material selection parameters required to achieve fully dense parts during direct Selective Laser Sintering of metal.Test MethQdsSeveral low melting point metals were tested to determine if they could produce full density SLS parts. Two nickel brazes, BNi-6 and BNi-7, were selected because of their relatively high melting temperatures of 875°C and 890°C.

Two nickel brazes, BNi-6 and BNi-7, were selected because of their relatively high melting temperatures of 875°C and 890°C. In addition, three tin based solders, 63Sn-37Pb, 50Sn-50Pb, and 96Sn-4Ag were selected because of their low melting temperatures. The nickel brazes were obtained from the Wall Colmonoy Corporation and carry their designation of Nicrobraz 10 for the BNi-6 and Nicrobraz 50 for the BNi-7. The composition of Nicrobraz 10 is llP-89Ni and the composition of Nicrobraz 50 is 14Cr-IOP-Ni balance.The rate of liquid infiltration into three dimensional porous media is known to depend on the liquid-vapor surface tension "flv' liquid viscosity n, and particle size d [I].

We measured "flv using one of two methods, the pendant drop for high melting temperature alloys and the sessile drop for low melting point alloys.The pendant drop weight method for determining surface tension was used for the high melting temperature alloys because no contamination from a substrate is added to the material being tested as the drop is formed homogeneously at the end of the rod. The liquid-vapor surface tension "flv can be determined by measuring the weight of a pendant drop after it has separated from the rod . As a drop forms at the end of a rod, the weight of the liquid drop is balanced by the surface tension. When the drop separates from the rod, some fraction of the liquid remains on the rod.

As a drop forms at the end of a rod, the weight of the liquid drop is balanced by the surface tension. When the drop separates from the rod, some fraction of the liquid remains on the rod. The surface tension can be calculated as:"flv = mo g F/Rlwhere mo is the mass of an "ideal" drop that completely separated from the rod, g is the gravitational constant, F is the correction to account for the fraction of the "ideal" drop that actually separates, and Rl is the radius of the solid rod. The factor F is related to R I/yl/3 (where Y is the volume of the drop after it has separated from the rod) and is empirically fitted to the equation:440(1)F = 0.14782 + 0.27896 (R1/yl/3) - 0.166(Rl/yl/3)2where Y is the volume of a drop and is calculated from the weight of a drop divided by the liquid density.To use the pendant drop method, the material being tested must be in rod form.

Since Wall Colmonoy Nicrobraz 10 and 50 are only available in powder form, rods of the material were vacuum cast. The powder was placed in a 3.0 mm I.D. glass tube with one end sealed, the tube was evacuated using a mechanical pump, and the other end of the tube was sealed using a propane torch. The tubes were placed in a furnace operating at 900°C and the powder melted and fused together into a rod. The glass simultaneously softened and, due to the pressure differential, collapsed around the liquid. After cooling, the glass sheath was removed from the metal rod.Figure 1 is a sketch of the apparatus used to measure the surface tension of different materials using the pendant drop technique. A rod of the material being tested was suspended from the bottom hook of the balance and the furnace was raised using a lab jack to heat the end the rod. As drops fell from the rod the weight indicated on the balance decreased by the weight of each drop.

As drops fell from the rod the weight indicated on the balance decreased by the weight of each drop. A computer program was written to interface with the balance and detect the weight change, average the drop weight, and calculate the surface tension.The low melting temperature alloys could be tested using a hot plate in a vacuum chamber. Glass microscope slides were used as the substrates for the sessile drop measurements. A 35 mm SLR camera with a 150 mm focal length lens and a 7x "close-up kit" lens assembly was used to photograph the drops.

A 35 mm SLR camera with a 150 mm focal length lens and a 7x "close-up kit" lens assembly was used to photograph the drops. The chamber was evacuated and back-filled with premixed 4% H2 in nitrogen.o Bottom 0 00 on ~ g Loading '----"==r=""----' Balance~RodL10000C FurnaceLab JackE:::::====:JFigure 2 is a sketch of a sessile drop defining the values of the drop diameter dm, the height of the drop from the maximum diameter to the top of the drop H, and the distance H', from the top of the drop to the intersection of the axis with a 45° tangent to the drop. There are two equations that can be used to measure the surface tension from the dimensions of a sessile drop, but only one is applicable to small drops where the top surface is not flat. The Dorsey equation was used to calculate the surface tension from the curvature of a sessile drop [2]. The Dorsey equation is:"flv = g P4 d~ r·Oi20 - 0.1227 + 0.0481 f)Figure 1.

The Dorsey equation was used to calculate the surface tension from the curvature of a sessile drop [2]. The Dorsey equation is:"flv = g P4 d~ r·Oi20 - 0.1227 + 0.0481 f)Figure 1. Pendant drop-weight apparatus for determining surface tensionwhere the Dorsey factor f is given by2H'f = dm0.4142The viscosity of the liquid solders was measured using a capillary flow tube. Figure 3 is a schematic drawing of the viscometer. The viscometer was constructed from an insulated graphite rod machined to form a reservoir above an alumina capillary tube. A quartz powder suspension in butanol was applied the exposed surfaces of the graphite to keep the graphite from oxidizing in the furnace. The quartz layer was not wet by the liquid metal and also prevented the liquid metal from reacting with the graphite. The viscometer was placed in a furnace and allowed to soak for one441(2)(3)(4)hour prior to testing.

The viscometer was placed in a furnace and allowed to soak for one441(2)(3)(4)hour prior to testing. The viscometer was then removed from the furnace and supported over a small collection vial of known volume. A tapered 2 mm graphite rod was used to plug the capillary tube while liquid metal was poured into the reservoir. The tapered rod was removed to allow molten metal to flow through the capillary. The time to fill the vial was recorded.The viscometer was calibrated using the known viscosity of liquid mercury and molten gallium metal. The equation used to calibrate the viscometer was:1ttr4 pnVp + 81t t L(5)where 11 is the viscosity of the liquid, r is the inner diameter of the capillary tube, P is the pressure head, t is the time, V is the volume of liquid passing through the tube in time t, L is the length of the tube, n is a calibration constant, and p is the density of the liquid.Figure 2.

Sketch of a sessile drop showing relationship of height and diameter used for the Dorsey equation (Equation 3) to calculate surface tensionThe infiltration rate of a liquid into a thin powder bed was measured by depositing a liquid onto the top surface of a thin powder layer. Water and n-butyl alcohol where deposited on 1.0 and 0.3 mm thick layers of 100 mesh and 325 mesh metal powders. After a relationship between spreading rate and surface tension, viscosity, powder size, drop volume, and layer thickness was determined, the spreading rate of molten solder in copper powder was measured to validate the relation for metals.To tie the infiltration relationship to the SLS process, a series of tests were conducted in which a C02 laser beam was focused at different locations on a powder bed of 80% BNi-6 and 20% 1018 steel for short periods of time.

Scanning mirrors directed the laser beam to a location, dwell on the spot for a short time and then jump to the next programmed location where the process was repeated with increasing dwell times. Three laser powers were tested: 10, 20, and 30 watts. The dwell time at each location was increased from zero seconds to 1.5 seconds by 0.25 second increments. After a series of sintered spots were created, a quick setting epoxy was applied to the bed and allowed to cure so the spots could be microsectioned. The infiltrated metal beneath the spots was examined using optical microscopy to determine the density and infiltration characteristics of the liquid into the metal powder bed under the laser spot.ResultsThe pendant drop experiments on BNi-6 indicated that the surface tension of this alloy in an atmosphere of premixed 96% N2' - 4% H2 with a residual of 1.7% oxygen from air, is 1360 mN/m at its melting temperature of 875 0 C.

The value of the surface tension of BNi-6 in 96% N2 - 4% H2 with 2.0% oxygen from air decreases to 1300 mN/m. When the residual oxygen was reduced to 0.7% by purging the glovebox with more premixed 96% N2 - 4% H2 the surface tension of the BNi-6 increased to 1580 mN/m. When the oxygen content was reduced further to 0.3% the surface tension increased again to a value of 1600 mN/m.Surface tension measurements using the pendant drop method were not possible on BNi-7 alloy because the outer surface of the rod formed a oxide film that prevented the formation of droplets at442the temperature (lOOO°C) and oxygen partial pressures (less than 0.1 % oxygen in 96% N2 - 4% H2) present in the glovebox furnace.3 em8 em4 emLThe surface tension of a liquid can be calculated at any temperature above the melting point (but below the boiling point) using the sessile drop test. However, this requires measuring the density of the liquid at the temperature of the surface tension test.

However, this requires measuring the density of the liquid at the temperature of the surface tension test. Figure 4 is a graph of the data collected on the densities of three Sessile drop solders at various temperatures. experiments indicated surface tensions of 50Sn-50Pb, 1.1 mm 10 Alumina 63Sn-37Pb, and 96Sn-4Ag solders to be 360, 370, and Capillary 440 mN/m, respectively. The literature value for the surface energy of 63Sn-37Pb in a vacuum is 492 mN/m [3]. The surface tension values reported here are lower than reported in the literature. This difference is perhaps due to the oxygen contamination in the chamber of about 0.1 %. Although an attempt was made to calculate the change in surface tension with changing temperature dy/dT, there was no statistically significant change of "(Iv in the temperature range of 200° to 350°C.Reservoir Graphite Body9.5 mm Riser TubeThermocouple3 ~ 18 cmReceiver~ OverflowFigure 3.

Viscometer used to determine viscosity of liquid metalsThe literature value for the d"{ldT of 63Sn-37Pb in a vacuum is 0.09 mN/m °C-l which is too small for us to detect in the data scatter. The surface tension data are presented graphically in Figure 5. The lack of a definite decrease in surface tension with increasing temperature is apparent.9.0 _....... .,- • • • • • Sn 50 • Sn 63 • Sn 96 .... •A model for the infiltration rate of a liquid into a thin powder layer was empirically developed from liquid spreading into a thin layer The test of metal powder. parameters were: layer thickness, the amount of liquid initially present, the particle size, the viscosity of the liquid, and the surface tension of the liquid. After data analysis, it was determined that the layer thickness and initial amount of liquid deposited on the top of the layer had no statistically significant impact on the radial rate of liquid infiltration.

The distance r the infiltration front advances in a radial direction is given by:......... .....8.8"'"'"'"'"'"'U 8.6 u 0,8.4 8.2the>- :!:: 8.0II); 7.8 C 7.6..•7.47.2400300450200350250(OC)Temperature...Figure 4. Density variations of solder with temperatureradius = 6.0.10-5 (time)0.5 (:vJO (diameter)-05where the radius of the advancing front is given in centimeters, time is in seconds, Ylv has units of mN/m, 11 has units of mPa·s, and the particle diameter has units of micrometers.443500(6)Data from the solder spreading in copper powder are graphed in Figure 6. The solid lines represent equation 6 with the corresponding viscosity, surface tension, and particle size values.

The solid lines represent equation 6 with the corresponding viscosity, surface tension, and particle size values. The data for Sn63 did not fit the equation due to the uneven radial spreading of the solder into the copper bed for that experiment.DiscussionE-z.Sn 50 • Sn 63 .Sn 96_ 500• • • • •• •• I • •Selective Laser Sintering of metal parts poses a challenge because both the processing conditions and the metal be controlled. The laser beam must be controlled so that only the low melting point particles melt. If there is too much liquid, created either by melting high melting temperature particles or by the formation of a liquid solution, the powder in the layer will tend to form spherical balls. It is critical that the liquid phase wet the solid phase completely and rapidly during the SLS process.

It is critical that the liquid phase wet the solid phase completely and rapidly during the SLS process. Since a liquid phase must be generated that wets the solid phase, both the solid and liquid phases must be selected so that their surface energies assure wetting and rapid capillary flow. The solubility of the solid phase into the liquid should be as low as possible consistent with rapid, complete wetting. The liquid phase must have a low viscosity so that the liquid can infiltrate the solid particles rapidly. The liquid phase should have a lower absorbance to the laser radiation than the solid so that the solid particles will be hotter than the liquid. This will promote capillary flow and rapid wetting while reducing the tendency of the liquid to form spheres.•E - 450 II: osystem must•• • • ••0) ; 400 I-•• •4D (,) ~ 350..:::IthetJ)3 0 0 +--+--+--+--+---1----+--+--+--+--+--+---+---+---+--I200350300250TemperatureFigure 5.

Surface tension data from sessile drop tests for solderWhether a liquid will wet a solid is determined by equilibrium thermodynamics. Young's equation [4,5] for wetting was developed from a mechanistic approach of balancing horizontal components of the three interfacial energies at equilibrium:"{sv - "{Is = "{Iv cosSwhere "{sv is the solid-vapor interfacial energy, "{Is is the liquid-solid interfacial energy, "{Iv is the liquid-vapor interfacial energy, and S is the contact angle the liquid makes with the solid. In both the pendant drop and sessile drop tests, the system was assumed to be at mechanical and chemical equilibrium; Le., the interface was static and there were no chemical reactions. Since the system was at or near equilibrium, neither the value of the contact angle nor the surface energy of the liquid changed with time.A liquid metal will tend to form droplets unless there is a solid that the liquid can wet and infiltrate into or spread over.

The infiltration model presented in equation 6 implies that the ratio of surface tension to viscosity 'YIvlTl should be as high as possible for the liquid to rapidly infiltrate the powder bed. The viscosity of liquid metals varies 'from about 0.2 mPa·s for alkali metals to around444(7)5 mPa·s for d-transition metals [6-8]. The surface tension of liquid metals encompasses the range of 1800 mN/m for d-transition metals to around 300 mN/m for the alkali metals and semi-metals [9]. A good low melting temperature metal for SLS processing would have high surface tension, but not higher than the high melting temperature metal, and a low viscosity. Since oxides reduce the surface tension and increase the viscosity of liquid metals, their formation must be avoided.2.01.81.6Sn 50 • Sn 96 Sn 6301.4-E u 1.2-III 1.0::s " 0.8 CIS a:0.60.40.20.002~'--'T"""im-e-(secl/2)34Figure 6.

Solder spreading in copper powderAnother condition for the metal SLS process to be successful is that the ratio of liquid to solid must be controlled so that there is sufficient liquid to form a continuous infiltration front to sweep porosity from the powder bed. Additionally, sufficient solid particles must be present to form a solid network that the liquid can wet. If the solid particles do not form a relatively rigid network, the liquid will form spheres with solid particles in suspension rather than infiltrating into the solid network. On the other hand, if there is insufficient liquid to completely displace the vapor, residual porosity will remain after the liquid solidifies. More liquid is required to displace the vapor than that calculated from simply replacing the volume fraction of vapor because the liquid must completely encapsulate the solid and form a high solids content suspension.

The amount of low melting point metal required is typically in excess of 65 volume percent [10,11], almost twice the amount of liquid required to merely replace the vapor fraction.The binary phase diagrams for nickel-phosphorous, iron-phosphorous, and nickel-iron show that there is a eutectic in both the iron-phosphorous and nickel-phosphorous system and that nickel and iron are completely miscible in each other. Because there is extensive solubility of the solid steel in the liquid braze, complete wetting is possible. The advantage of using the nickel braze rather than tin-lead solders or copper-based brazes is that there appears to be less oxide film on the nickel-phosphorous particles Gudging from the lack of dross on the SLS layers), so wetting occurs very rapidly and completely.

The same behavior can be achieved with solder if the mixed solder and copper powder are fluxed with a very aggressive flux (Alpha 90 or Kester SP-30).Figure 7 shows the top surface of a 80% BNi-6 - 20% 1018 steel powder mixture that has been fused with a 30 watt C02 laser beam. The surface is concave due to slumping of the powder layer when the liquid infiltrated the powder bed. The BNi-6 has completed displaced the vapor phase and it appears to be fully dense. The extent of infiltration by the liquid is marked by loose powder445around the perimeter of the part.The binary phase diagrams for copper-tin, copper-lead, and tin-lead indicate that there are no common eutectics between the three systems, unlike the iron-nickel-phosphorous system; instead, there are several intermetallic compounds. The formation of intermetallic compounds may increase the viscosity of the molten solder and prevent it from infiltrating the pores of the powder bed [12].200 IlmFigure 7.

The formation of intermetallic compounds may increase the viscosity of the molten solder and prevent it from infiltrating the pores of the powder bed [12].200 IlmFigure 7. SEM micrograph of a 80% BNi-6 - 20% 1018 steel powder mixture that has been fused with a laser beamThe literature value for the viscosity of pure nickel is 5.2 mPa·s [6,7,13], and the Ni-P eutectic probably has a viscosity much lower than the pure element. Copper has a viscosity of 3.36 mPa·s and a surface tension of 1285 mN/m [6,7,13]. Therefore, the 'YlviTl parameter for copper is 382 and it should infiltrate steel powder in a similar fashion as the BNi-6 braze alloy. However, the reflectivity of atomized copper powder (as opposed to electrolytic powder) at a C02 laser wavelength of 10.6 Ilm is much higher than the reflectivity of nickel or iron [14].

However, the reflectivity of atomized copper powder (as opposed to electrolytic powder) at a C02 laser wavelength of 10.6 Ilm is much higher than the reflectivity of nickel or iron [14]. For the same laser energy density, much less copper is melted, so there is less liquid formed to infiltrate the pores between the iron particles.ConclusionsDuring direct metal SLS, the metal powder layer consists of low melting point particles, high melting point particles and vapor. The rate limiting factor for how fast a laser beam can scan across the metal powder bed and leave behind a fully dense area is related to the sum of the time required to melt the low melting point particles to create a liquid and the infIltration rate of the liquid metal into the powder.

The vapor fraction of a powder bed must be expelled by an advancing liquid front if the gas is to be removed since diffusion of the gas through the liquid or bubbles rising through the liquid (Stoke's Law) require a longer time frame than available during Selective Laser Sintering (SLS). However, if sufficient liquid phase is present so that the front is continuous, the vapor can be displaced by the advancing liquid and the area of the metal powder bed scanned by the laser will be fully dense.446The infiltration radius r, of a liquid metal into a porous metal powder layer is given by the equation:radIus =6.0.10-5 (tlme)O.5 -:n..(YIV)2.0.(diametertO.5where the radius of the advancing front is given in centimeters, time is in seconds, Ylv has units of mN/m, 11 has units of mN/m, and the particle diameter has units of micrometers. From this equation we can draw several conclusions regarding selecting a metal powder for the SLS process.

From this equation we can draw several conclusions regarding selecting a metal powder for the SLS process. The surface tension Ylv should be as high a possible consistent with Young's equation regarding wetting of the solid. The viscosity of the liquid should be as low as possible and the particle size of the solid metal should be as small as possible.If metal powders are selected with the appropriate surface tension, viscosity, wetting contact angle, and oxidation resistance and mixed in the appropriate ratios, it is possible to achieve fully dense SLS parts.AcknowledgementsThis study was made possible by support from a State of Texas Advanced Research Project, ATP 116: Selective Laser Sintering: Direct Metal Fabrication and ONR grant#NOOO14-92-J-1314.References1.Semlak, K.A., & Rhines, R.N.: "The rate of infiltration of metals" Trans Met Soc AIME 212 (6) 325-331 (1958) Iida, T. & Guthrie, R.I.L.: The physical properties of liquid metals Clarendon Oxford (1988)3.

ndamentals of Liquid Phase Sintering Related to Selective Laser SinteringD.E. Bunnell, D.L. Bourell, and H.L. MarcusThe Center for Materials Science University of Texas at AustinAbstractLiquid phase sintering is one of the underlying principles that must be modeled and understood when the Selective Laser Sintering (SLS) process is used. This paper describes the initial studies being conducted to measure surface tension of metal alloys used for SLS. A low melting point solder was used to verify the wetting balance and pendant drop techniques and equipment liquid-vapor, and solid for determining surface tension. vapor surface tension of 80 Sn - 20 Pb solder on mild steel was determined to be 245, 417, and 662 dynes/cm.The liquid-solid,Introduction Selective Laser Sintering (SLS) is a process that uses a rastering laser to sinter powder particles into a computer defined shape.

Because the laser is typically rastered at speeds of 1 to 10 cm per second, If the laser spot may only be at sintering temperatures for a few seconds. particles are to be sintered during this short period, either melting or a reaction must occur to form particle necks. Liquid phase sintering chemical (LPS) differs from conventional sintering in that a liquid phase allows for particle rearrangement and contact dissolution which accelerates the initial shrinkage rate and is the underlying principle of SLS processing of two phase (low and high melting point alloys) systems.the particles under theTo evaluate the phenomena that control LPS during SLS, several considered: surface tension of liquid-solid and liquid-vapor interfaces, rate of the solid phase in the liquid phase, powder bed as the laser rasters over it, modeling of the initial particle rearrangement during liquid phase sintering, and applying the model fabricate a multilayer metal part using SLS.

Results of the initial studies of these factors will be presented here.factors must bethe rate with which a molten metal will wet a solid metal,thethe dissolution the temperature transition of atoLiquid Phase Sintering TheoryKingery proposed three stages for liquid phase sintering 1 is the formation of a liquid phase which flows and facilitates the rearrangement of particles. dissolves the solids and the compact density increases by solution and reprecipitation. compact become homogeneous and the melting point of the liquid exceeds the processing temperature.The first stageThe second stage occurs when the liquid phaseThe final phase is the formation of a solid skeleton as theresults in The first stage of liquid phase sintering, particle rearrangement, the most rapid and largest extent of densification 2 the only ,3.

During SLS, part of the powder bed that contains liquid is the small area directly under the moving laser beam and the small area trailing the current position that379has not cooled below the melting point of the liquid phase. These areas will only experience a molten second phase for a short period of time, of the order of seconds, before being cooled to the temperature of the surrounding solid material. Therefore, it behooves us to model particle rearrangement kinetics so as to characterize SLS.the liquid phase sinteringDuring the initial stage of LPS the mechanism for densification is particle rearrangement driven by surface tension forces 4 . Kingery progosed that the densification of the compact could be modeled by the equationWhere V is the volume of the compact, en is the coordination number, Yl v the liquid-vapor surface tension, t is the melt viscosity.

vapor surface tension, and ~, the liquid viscosity.isis time,r is the particle radius, and ~ the liquidFrom this equation, we need to find Yl v 'The driving force for liquid phase sintering is the capillary force that a wetting liquid exerts between two solids 6 The capillary force consists of two components, surface tension and a pressure difference due to the meniscus curvature. it is necessary to find Yl s and Ysv ',7.To determine the capillary forces, the liquid-solid and solid-vapor surface tensions.The kinetics of wetting are not completely understood. preceding the advancing liquid front removal are complex and involve diffusion of metal vapors through pores in the oxide and changes in surface tension due to surface active elements.

preceding the advancing liquid front removal are complex and involve diffusion of metal vapors through pores in the oxide and changes in surface tension due to surface active elements. Since wetting of metal particles is required for liquid phase sintering during SLS, the kinetics of wetting need to be determined to modelThe reactionsthat are responsible for oxide layerthis process.The wetting balance test provides information on the transient wetting characteristics of a liquid-solid couple8 . Figure 1 is a sketch of the three stages that a solid sample experiences during a wetting balance test. During the initial wet and is lifted by the surface tension of the liquid. Within a fraction of a second, the liquid starts to react with the surface layer and begins to wet the exposed surface of the solid.

Within a fraction of a second, the liquid starts to react with the surface layer and begins to wet the exposed surface of the solid. meniscus has formed.immersion of the solid into the molten metal,the sample is notThe third stage is when the equilibriumInitial PenetrationPortial WettingComple te Wettil"lgFigure 1. Stages of wetting during wetting balance test.380The value of Y (the liquid-solid surface tension) can be determined from a force balance that perimeter of the sample, buoyancy term related to the volume of liquid displaced:includes three terms,the weight of the meniscus around thethe surface tension that pulls the liquid up, and aLF y = 0 = (L\wt'g)/lYl s / cose + I1p'VdWhere L\wt constant, 1 is the perimeter of the sample, e is the contact angle of the liquid-solid interface, liquid, and Vd is the volume of liquid displaced.

If the sample penetration into the liquid is small, neglected and from the force balance, Yl sis the weightincrease of the sample, g is the gravitationalI1p is the difference in density between the solid andthe buoyancy term can beis given by the relationship:Yl s = (l1wt'g)/l. coseThe liquid-vapor surface tension (Yl weight of a pendant drop9. Figure 2 1S a sketch of a pendant drop showing the two curvatures. As a drop forms at the end of a rod, two curvatures control when it will separate. curvatures of the pendant drop are related to the surface tension by the Laplace equation'o:) can be determined by measuring the XIf the drop is small, gravity may be neglected and theI1P = Yl)Hwhere 2H is the Gaussian average curvature and I1P is the pressure difference.

For a symmetrical drop where both curves are revolutions about an axis, average curvature is given by:the2H = (l/R, + l/Rz) In practice, an empirical relationship exists between the diameter of the rod and 2H.The surface tension can be calculated as:Yl v = m·g·F/R,where m is the mass of the drop, ~ is the gravitational constant, F is the correlation factor related to r/V /3 (where V is the volume of the drop after it has separated from the rod), and R,is the radius of the solid rod.::==:'1I:=~;=:l/. R7.)~~/ (Figure 2. Curvature on a pendant drop381(1)(2)The factor "F" is empirically fitted to the equation:F = 0.14782 + 0.27896(rjV'/3)0.166(rjV'/3)2(3)Where V is the volume of a drop.To determine the solid-vapor surface tension, Y , solidified metal was measured.

At equilibrium,Sthe condition for a liquid to wet a solid is given by Young's equation":the contact angle of the(4)Since Yl s and Yl v are measured using the wetting balance and the pendant drop tests,to find Ysv allthatis needed is the contact angle 8.EquipmentA "wetting balance" was used to measure the rate of wetting and to provide data used to calculate Yl s ' The wetting balance consists of a bottom loading balance accurate to 0.01 gram with a computer interface, a 1000°C top-loading furnace with a graphite crucible, and a lab jack that raises and lowers the furnace. solid is suspended from the bottom of the balance and the liquid metal melted in the furnace below the balance. When the solid sample and the molten metal are at activated to record the weight changes and the furnace is raised using the lab the sample touches the liquid surface.

When the solid sample and the molten metal are at activated to record the weight changes and the furnace is raised using the lab the sample touches the liquid surface. At this time, an electrical jack until circuit from the balance to the crucible is closed that deactivates the lab jack motor. This procedure allows for a reproducible (and small) penetration depth of the solid sample into the liquid so that buoyancy effects can be minimized. After 10 seconds the furnace was lowered and the sample was sectioned and metallographically examined to determine the thickness change.the liquid-solid surface tension.In practice, a sample of theisthe correct temperature for the test,the computer interface isInitial melting point materials. Mild steel sheet was the solid sample and the 80Sn 20Pb solder was used as the molten metal.

Mild steel sheet was the solid sample and the 80Sn 20Pb solder was used as the molten metal. This alloy was chosen since it wets steel and melts at a temperature below 300° C.tests were performed to test the suitability of the methods on lowThe same equipment was use to measure the liquid-vapor surface tension (Yl v ) using the drop weight method. A rod of the alloy was suspended vertically from the balance and the furnace was heated to 300°C for the solder alloy. The furnace was raised until melting point and liquid drops formed and fell weight change of the rod, which corresponds to the weight of the drop, was recorded for several drops and from these data, Yl v was calculated.just the end of the brazing rod was heated to itsinto the empty crucible.TheRESULTS &DISCUSSIONThe data from a wetting balance test of steel and Sn80 solder at 300°C are presented in Figure 3. was a constant 1.82 grams as the molten solder was being raised to contact edge of the sample.

was a constant 1.82 grams as the molten solder was being raised to contact edge of the sample. contacted the liquid solder and pushed down on the surface. There is no initial wetting and the steel solder.For the first three seconds the weight of steel sample the The weight decreased rapidly at 3.19 seconds as the steelformed a depression in the surface of the moltenThe rate that the weight decreased is related to the speed at which382the solder pot it penetrates the surface of the molten solder does not reveal any surface tension data because it is dependent on many parameters such as the velocity of the solder pot and the shape of the solid sample edge.is raised and the rate of wetting.The weight of the sample as20/80 solder on steel 300 C, Alpha 90 Flux3.52.5f/)§ 2 (;) -----~1.50.51.653.35.658.3~)10111;513.291-1,94SecondsFigure 3.

Wetting balance data for 80Sn-20Pb solderThe motor lifting the solder pot stopped when an electrical contact was made between the solder and sample, but millimeter more after contact. Between 3.2 and 5.04 seconds, the solder wet the sample. During this time the molten metal reacted with any surface films on the steel and wet the solid. After five seconds the equilibrium meniscus formed and the weight the solder pot was lowered at fifteen seconds. There is a spike in the weight of the sample as the liquid formed an inverted meniscus. spike is not relevant because the sample is in motion and equilibrium conditions do not apply.inertia allowed the potto rise aleveled off at 3.37 grams and remains there untilThe height of thisThe plot of weight versus time shows that the wetting time is approximately 2.2 seconds. This time represents the time from when the sheet of steel touched the molten solder surface to the time when the solder meniscus stopped advancing up the sheet.

This time represents the time from when the sheet of steel touched the molten solder surface to the time when the solder meniscus stopped advancing up the sheet. rising above the surface of the solder is 1.55 grams. The wetting angle was measured as 0° using an optical microscope at 100x so the value of Yl s was calculated using equation 1 to be 245 dyne/em.The change in weight due to the solder meniscusFigure 4 is a plot of sample weight versus drop number during a pendant drop test. dipped in Kester SP-30 solder paste flux was used.In this test a 80Sn-20Pb solder rod with a diameter of 0.32 em and The first drop weighed3830.15 grams and was most likely a drop of flux. All weighed between 0.25 and 0.27 grams.

All weighed between 0.25 and 0.27 grams. From these data, F was calculated from Equation 3 to be 0.2368, and the liquid-flux surface tension was calculated using Equation 2 to be 417 dyne/cm.the remalnlng dropsThis compares with a literature value of 450 dyne/cm 12 but since they were measuring wetting in an inert gas while we were using a zinc chloride based flux, our value should be somewhat lower than the reported value. test where no flux was used and the sample was surrounded by air, Yl v was determined to be 662 dyne/cm.In a second the value of65\J 4 o 0:: .... o 3 :c .Q'> (I) 5 280 n0 rJb Solder flux.. 0.27 0.27325 C, ZnCII I III0.150.260.270.260.25o2 345 Drop Number67Figure 4. Data from pendant drop test on 80Sn-20Pb solder with fluxWithout flux, in order to determine the solid-vapor surface tension from Young's equation the liquid-vapor surface tension obtained using flux was used for this experiment.

Because of this, the solid-vapor surface tension is actually the solid-flux surface tension. The contact angle 8 was measured from the cross section of the steel, and Ysv was calculated using equation 4 to be 565 dyne/cm.the solder will not wet steel. Therefore,SummaryThese techniques for measuring surface tension (y) and kinetic data are the initial particle rearrangement stage of liquid phase necessary to model sintering. The pendant drop technique is an accurate and reliable method of determining Yl and the wetting balance test gives not only Ylf' but wetting kinetics data as well.

The pendant drop technique is an accurate and reliable method of determining Yl and the wetting balance test gives not only Ylf' but wetting kinetics data as well. the particle rearrangement process that occurs as the first stage in LPS.By using this information we hope to De able to model In384turn, three dimensional parts.the model of LPS will be used to optimize SLS of metal powders to makeThis work is preliminary, Future work will include the surface temperature measurement of the brazing alloy in a reducing atmosphere.to test the equipment and methods.AcknowledgementsThis study was made possible by support from a State of Texas Advanced Research Project, ATP-116: Selective Laser Sintering: Direct Metal Fabrication and ONR grant #N00014-92-J-1314.References1. Kingery, W.O. "Densification during Sintering in the Presence of a Liquid Phase" J. Applied Physics 30 3, 301-306 (1959)2. Gessinger, G.H., Fischmeister, H.F., and Lukas, H.L.; "A Model2.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFundamentals of Polymer Crystallization in Laser Powder Bed Fusion for New Material ScreeningCamden A. Chathama, Samantha J. TalleybaAdvanced Engineering Division, Savannah River National Laboratory, Savannah River Site, Aiken, 29808, SC, USA bMaterials Engineering, Kansas City National Security Campus, 14520 Botts Rd, Kansas City, 64147, MO, USAAbstractAlthough laser powder bed fusion (PBF/LB) was one of the first industrially viable additive manufacturing (AM) methods for end-use part production, polyamides remain grossly dominant at both the commercial- and research scale. The research community continues to develop and refine “rapid screening” methods for evaluating the suitability of new polymers for PBF/LB.

The research community continues to develop and refine “rapid screening” methods for evaluating the suitability of new polymers for PBF/LB. The so-called “SLS Process Window,” which is the difference between melting and crystallization temperature measured at 10 K min- as originally outlined in the patent literature, is perhaps the most often reported screening method. Although perhaps appropriate as part of a larger study, the 1 simplistic guidelines put forth by the “SLS Process Window” are not sufficiently scientifically rigorous to understand how crystallization kinetics affects successful 3D printing. The common understanding of the SLS Process Window omits details from published theories of polymer crystallization. as evidenced by published assumptions and methods in PBF/LB process modeling papers.

as evidenced by published assumptions and methods in PBF/LB process modeling papers. The authors explain polymer crystallization in the PBF/LB context and propose replacing the “process window” with crystallization halftime and physical gelation for new material screening. These measurements better represent behavior critical for ensuring a lengthy coexistence of solid powder and molten polymer affecting warp-free parts.IntroductionIn general, thermoplastic additive manufacturing (AM) can be divided into material extrusion (MEX) for amorphous polymers and laser powder bed fusion (PBF/LB) (originally commercialized as Selective Laser Sintering; SLS [1]) for semi-crystalline polymers. All thermoplastic manufacturing methods require in-depth understanding of the relationship between polymer viscosity, temperature, and pressure. For semi-crystalline polymers, this includes the first- order thermal transitions for melting and crystallization, which have a profound effect on viscosity.

For semi-crystalline polymers, this includes the first- order thermal transitions for melting and crystallization, which have a profound effect on viscosity. The rapid heating and cooling cycles arising from PBF/LB’s distinctive laser scanning is a complex environment for melting and crystallization often considered to be quasi-isothermal by the modeling community [2]. Despite the complex temperature profiles, many investigations into adapting or evaluating the potential printability of polymers in PBF/LB hinge on the so-called “SLS Processing Window.” Constructing the SLS Processing Window involves running a dynamic and observing the differential scanning calorimetry (DSC) experiment at either 10 or 20 °C min degree of separation on the temperature axis between melting and crystallization behavior.

Statements published in the literature regarding this technique indicate that (i) a wider window is−185better [3, 4] and (ii) the observed separation of melting and crystallization behaviors must occur for successful printing [3, 5]. Other authors claim observing this behavior indicates that (iii) the scanned polymer remains molten throughout the entire build [6, 7, 8]. The authors believe that, although these statements originated from scientifically sound claims from the patent literature, the common use in published literature falls short of accurately describing polymer behavior in a manner consistent with the historic body of polymer crystallization literature. Failing to recognize the simplifications made in the patent literature is hindering the research community from developing and/or adapting new polymers for PBF/LB.

Failing to recognize the simplifications made in the patent literature is hindering the research community from developing and/or adapting new polymers for PBF/LB. In this paper, the authors present the original usage and merits of the PBF/LB Processing Window as a cursory first glance at understanding the relationship between viscosity, morphology, and temperature before highlighting the simplifications made for the sake of patent claims and expounding on the fundamentals of polymer crystallization applied to the PBF/LB context.Figure 1: The oft reported “SLS Processing Window” overlay of DSC heating and cooling data collected at 10 °C min-1 for two AM-grade nylon-12 materials and a pellet nylon-12 control.

Annotations highlight the separation between melting and crystallization behavior.The Historical Use of the PBF/LB Processing WindowThe technology known as “laser based polymer powder bed fusion (PBF/LB-P)” according to ASTM/ISO 52900-21 [9] was originally patented and commercialized as “Selective Laser Sintering (SLS)” by Carl Deckard and DTM Corporation [10, 1]. The method describes using a laser beam to selectively heat and fuse together powdered polymer in a layerwise manner to generate a 3D object. The so-called “SLS Process Window” originates from subsequent DTM Patent literature. Follow-on patents in 1994 and 1997 by DTM describe commonalities among materials that could be successfully printed by this method. [11, 12].

Follow-on patents in 1994 and 1997 by DTM describe commonalities among materials that could be successfully printed by this method. [11, 12]. The patents describe either “non-overlapping” (1994) or “slightly overlapping” (1997) temperature regions for observed melting and crystallization behavior when the material is evaluated by a dynamic DSC experiment measured using a reasonable heating rate between 10-20 °C min- . The patents describe example successful materials being printable with as much as a 13 °C temperature overlap accounting for 1 6.2 % of the integrated area underneath the melting and crystallization curves. Conversely, negative examples of unsuccessfully printed materials are showing a 24 °C and 21 % overlap in melting and crystallization behavior [12].

Conversely, negative examples of unsuccessfully printed materials are showing a 24 °C and 21 % overlap in melting and crystallization behavior [12]. The patent goes on to state that although this method is not a measurement of crystallization rate, it is related to crystallization rate.The vague patent language has apparently influenced both the reported figures and the tone of discussion in the scientific literature. Interestingly, scientific articles often provide less discussion than the patents on the fundamental physics principles underlying the “sintering window.” Some simply report the separation in melting and crystallization temperature without86mention of the rationale that this measurement is related to crystallization rate [13].

Others explicitly state that the observed separation of behavior is critical for successful printing [3, 5, 7] even though this is counter to the 1997 patent, which allows for small overlaps [12].Bain’s 2019 summary of the “so-called sintering window” asserts the intended goal of the measurement as demonstrating a slow crystallization rate to afford both thoroughly coalesced particles and no out-of-plane warping [14]. “Slow” is a vague and relative term that is not particularly helpful in PBF/LB material screening. “Slow” is not quantitative; one cannot know by this method only if it is “slow enough” for the PBF/LB context.Warping is another aspect of DTM’s patents that is rarely discussed in scientific literature. They make the claim outlining a critical minimum storage modulus for the polymer being printed that will not allow more than 50 µm of out of plane curling.

They make the claim outlining a critical minimum storage modulus for the polymer being printed that will not allow more than 50 µm of out of plane curling. Crystallization and its accompanying discrete reduction in specific volume is often cited as the major cause of out of plane warping leading to build failure; however, little discussion is given to the forces opposing warping including in review papers like those written by Bain [14] and Chatham, et al. [15].Fundamentals of Polymer CrystallizationThis section covers selected aspects of current understanding in the polymer science community of polymer melting and crystallization phenomena at a high level. References in this section are not explicitly tied to the PBF/LB or any manufacturing context. It is included to orient the reader around the wider wealth of research reported from studying polymer melting and crystallization.Equilibrium melting and supercooling.

It is included to orient the reader around the wider wealth of research reported from studying polymer melting and crystallization.Equilibrium melting and supercooling. The equilibrium melting temperature (Tm ) is the temperature above which a given polymer will never form stable regions of crystallinity. It 0 designates the threshold where melting and crystallization are in equilibrium for an extended-chain crystal (i.e., no folds in polymer chain conformation) [16]. It is often calculated based on Hoffman- Weeks theory extrapolating a linear fit of measure peak melting temperatures following complete crystallization at an independently determined isothermal crystallization temperature [17].

The difference between Tm and each independently determined (i.e., set by the researcher) temperature is termed the “undercooling” or “supercooling.” When PBF/LB - focused papers or patents employ 0 the term “supercooling” they are more often referring to the observed onset of crystallization and its difference from the observed onset of melting, which is a very different phenomenon.Crystallization. As the theory states, a polymer will crystallize at each and every temperature less than Tm . The question is not “will this polymer crystallize at this temperature?” but rather “how long will it take for this polymer to crystallize?” Crystallization comprises both 0 nucleation and growth. Depending on the chosen undercooling, the rate of crystallization will be governed by either the system’s ability to nucleate (if closer to Tm ) or the available energy in the system for chains to move into place and grow existing crystals (if closer to Tg).

The rate of 0 crystallization follows a parabolic curve and reaches its maximum somewhere between Tg and . Therefore, changes in temperature close to the maximum rate of crystallization have a smaller impact on rate of crystallization than temperature changes closer to the extremes of Tg and Tm . It is common to measure crystallization halftime (i.e., the time to reach 50% crystallized at a given 0 temperature by DSC) as the overall rate of crystallization. Alternatively, crystallization rate can be determined by monitoring the spherulitic growth rate from laser light scattering or polarized optical microscopy.87Relationship between crystallization and arresting diffusive motion (physical gelation). It is important to note that while segments of chains are incorporated into stable crystalline regions, the entire chain is not necessarily included in the crystal.

It is important to note that while segments of chains are incorporated into stable crystalline regions, the entire chain is not necessarily included in the crystal. Real crystals generally are observed to follow either adjacent re-entry or switchboard models of chain-folded lamella instead of extended chain arrangements. As regions of crystallinity develop in a polymer, they restrict the mobility of the pieces of polymer chains remaining in the amorphous bulk, thus increasing the viscosity. It is also important to note that a polymer may not be “100 % crystalline” when it is deemed “100 % crystalized.” A polymer may be considered highly crystalline with 40-60 % v/v regions of crystallinity. The key attribute is that due to the long chain nature of common thermoplastic polymers, the viscosity has become prohibitively high for long-range diffusive motion to occur, thus kinetically preventing molecules from reaching the crystal front.

In this way, it is possible and likely to have different absolute values of crystalline content (vol. %) at the “50 % crystallized” point in time to determine crystallization halftime.The relationship between crystallization and viscosity results in the phenomenon known as physical gelation. At this critical volume fraction of crystalline regions, the polymer behaves like a gel and long-range molecular motion is arrested. The point of physical gelation has been measured in literature using dynamic mechanical analysis (DMA). Using DMA, one can determine the temperature where measured tan(δ) is independent of frequency, such as reported by Hirayama, et al. [18].

Using DMA, one can determine the temperature where measured tan(δ) is independent of frequency, such as reported by Hirayama, et al. [18]. This temperature is taken as Tgel and can be mapped to a critical percent crystallinity through DSC using the matching rate of cooling.RoadmapAccording to Google Scholar, the article in which the following quote appears has beencited 1636 times at the date of this manuscript.“As the molten polymer cools down below Tm, polymer crystals nucleate and grow, recreating regions of ordered molecular chains (crystallites) mixed up with disordered amorphous regions...However, the freezing of the polymer at Tm coincides with an important shrinkage (phenomenon not occurring with amorphous polymers), that may induce geometrical inaccuracies and distortion of the part. A good way to prevent this is to preheat the polymer powder to a temperature slightly below its melting temperature and keep it there for a certain time after consolidation.

A good way to prevent this is to preheat the polymer powder to a temperature slightly below its melting temperature and keep it there for a certain time after consolidation. [19]”This is a good example of the dangers of imprecise language. So long as one understands 0),” there “melting temperature” in the quoted text to mean “equilibrium melting temperature (Tm is no issue. But if one incorrectly assumed “Tm” to be the peak melting temperature observed from a dynamic DSC experiment, that person would have an incorrect understanding of the physics governing PBF/LB-P.This example misconception is one the authors aim to remedy or at least raise awareness.

For the remainder of the manuscript, the authors discuss additional experimental methods that more thoroughly capture the interplay between temperature, viscosity, and crystallization in the PBF/LB context than the “sintering window.” This includes reporting the equilibrium melting temperature, crystallization halftimes, and physical gelation for two commercially available nylon- 12 powders marketed for PBF/LB and a generic nylon-12 control. For the sake of clarity, the authors use the distinguishing language “polymeric material” when referring to total material88formulations including not only the base polymer but also processing aids and additives typical of commercial materials.Materials and MethodsTwo commercially available nylon-12 powders were studied for this work. Duraform 2200 laser systems (λ = 10.6 is a common nylon-12 powder that is white in color sold for use with CO µm).

Duraform 2200 laser systems (λ = 10.6 is a common nylon-12 powder that is white in color sold for use with CO µm). Sintratec PA12 is a black nylon12 powder sold for use with shorter wavelength laser systems (e.g., λ = 1065 nm, 440 nm). A nylon-12 pellet-form sample purchased from Sigma Aldrich was chosen as a “non-AM grade” control material. This pellet is optically clear and does not contain additional compounds for any particular form of manufacturing.2Differential scanning calorimetry (DSC) was performed using a Mettler Toledo DSC 3+. All samples were run by typical heat-cool-heat experimental method at 10 °C min-1. These were combined with thermogravimetric analysis (TGA) experiments on a Mettler Toledo TGA 2 at 10 °C min-1 to construct the typical “SLS sintering window.” These experiments were conducted under nitrogen atmospheres.Isothermal crystallization experiments were carried out on the same DSC 3+ to determine equilibrium melt behavior.

Samples were heated above 200 °C for 10 min before quenching to the isothermal temperature and holding for 130 min. The Sigma Pellet and DF2200 samples were held at 166, 164, 162, 160, and 158 °C whereas the Sintratec PA-12 material was held at 161, 160, 159, 158, and 157 °C to obtain suitable data.Physical gelation experiments were conducted on a Mettler Toledo DMA/SDTA 1+ using liquid nitrogen cooling and a special shear fixture designed for low viscosity samples. Samples were prepared by melting raw material (powder or pellets) into an approximately 1.5 mm thick rectangle to ensure a homogeneous sample during testing. These rectangles were loaded into the DMA at room temperature. Samples were held in the melt state at 200 °C for 5 min before cooling at 1 °C min-1 to 130 °C. Modulus and phase angle were collected simultaneously using a frequency series of 1, 5, 10, 100, and 150 Hz. The Mettler Toledo DMA/SDTA 1+ switches between operating in a stress- or strain-controlled manner.

The Mettler Toledo DMA/SDTA 1+ switches between operating in a stress- or strain-controlled manner. The limits of 30 N force amplitude and 100 µm displacement amplitude were set in the method. One cycle is defined by satisfying a logical “OR” condition of either the force amplitude or displacement amplitude. Samples were observed to satisfy the 100 µm displacement condition before reaching 30 N over the entire range of temperatures tested.Results and DiscussionObserved Melting and Equilibrium MeltingDistinguishing between “observed melting behavior” and the “intrinsic equilibrium melting temperature” is critical for understanding how semicrystalline polymers behave during thermal processing. When discussing the Process Window, it is often said that the powder bed surface temperature is set slightly below the onset of melting [15]. The more precise statement is that it is set just below the observed onset of melting of the powdered feedstock.

The more precise statement is that it is set just below the observed onset of melting of the powdered feedstock. This observed onset of melting is not intrinsic to a given polymer, but rather is a manifestation of a specific material formulation previously processed at a set of conditions to elicit a chosen morphology. Material formulation, which includes not only the chemical identity, molecular weight, and89molecular weight distribution of the base polymer (e.g., polyamide-12) but also the various commercial “processing aids” (e.g., viscosity modifiers and crystallization packages) and “value adding fillers” (e.g., colorants and reinforcing fibers), affects both the observed and equilibrium melting behavior as different formulations can influence the type of crystalline order (i.e., a different lattice structure or inter-chain spacing).

Figure 2 shows the extrapolation from a linear fit of plotting peak melting temperature against isothermal crystallization temperature for DF2200 (white nylon-12), Sintratec PA12 (black nylon-12), and a nylon-12 pellet from Sigma Aldrich used as a non-AM grade control. The intersection from linear extrapolation and the line where Tm = Tc is taken to be Tm by the Hoffman-Weeks approach. The calculated spread of equilibrium melting temperatures is nearly 7 °C with the pellet control presenting the highest equilibrium melting temperature.0The basic definition of equilibrium melting temperature is the theoretical melting temperature for a perfect extended chain crystal [16]; it is therefore independent of processing conditions. Both observed and equilibrium melting behavior are important for understanding how a given polymeric material behaves during PBF/LB.

Both observed and equilibrium melting behavior are important for understanding how a given polymeric material behaves during PBF/LB. Equilibrium melting temperature is key to understand for mapping the overall possible process parameter space while observed melting behavior helps to contextualize a specific instance of prior processing. For example, the observed melting behavior of nylon-12 [20] and poly(phenylene [21] change with repeated processing; however this will not affect Tm as prior processing does not alter fundamental structure. Processing any given material imparts a specific formulation morphology, which includes crystal size, crystal size distribution, and the ratio. crystal-to-amorphous These factors may greatly influence the observed melting behavior without altering Tm.

Equilibrium melting temperature is calculated to be 194.3 °C for the control pellet, 191.3 °C for Sintratec PA12, and 187.2 °C for DF2200.2:morphology within a polymer without altering the equilibrium melting behavior.Onset of melting observed in the first heat of a DSC experiment is apt and suitable for determining the absolute upper-bound for bed temperature, as this experiment adequately captures the morphology induced by the powder making process (for virgin powder) or the morphology induced by repeated use in PBF/LB.900Table 1: Tabulated Results from DSCTm, onset Tm, endset Tm, peak HmTc, onsetSample[°C][°C][°C][J g-1][°C]DF2200171.7181.3171.621.2147.7Sintratec PA12173.1180.1177.724.6157.4Sigma Pellet (Control)162.8182.9177.751.9152.8Work from Zarringhalam, et al. highlighted the core/shell partial melting phenomenon they observed following PBF/LB printing of nylon-12 [22].

highlighted the core/shell partial melting phenomenon they observed following PBF/LB printing of nylon-12 [22]. This sort of behavior is expected if the entire particle does not reach the equilibrium melt temperature; only above Tm will all crystalline order surely be erased. Therefore, the authors suggest Tm and not simply Tm,end as the more ideal 0 minimum threshold temperature resulting from laser scanning. 0Crystallization KineticsGiven that the equilibrium melting temperature is the threshold below which crystallizable polymers will crystallize, the next question becomes “how long will it take to crystallize?” Predominant work mathematically describing the nucleation and growth of polymer crystallization was published by Lauritzen and Hoffman [17] and Avrami [23]. Many researchers use Equation 1 from Lauritzen and Hoffman or Equation 2 from Avrami as the starting point for studying a particular polymeric material in a particular processing context.

Many researchers use Equation 1 from Lauritzen and Hoffman or Equation 2 from Avrami as the starting point for studying a particular polymeric material in a particular processing context. Either isothermal (Lauritzen- Hoffman) or non-isothermal (Avrami) experiments can be conducted and fit to these equations for extrapolating beyond tested conditions or feeding predictive models.XctexpKtn() = 1 −(−)Figure 3 depicts the measured crystallization halftimes and mathematically fit data for the two PBF/LB nylon-12 powders and pellet control material. The figure focuses on the temperature region of interest for typical PBF/LB bed temperatures; however, the full shape of the data is a “U” nestled between Tg and Tm.

The figure focuses on the temperature region of interest for typical PBF/LB bed temperatures; however, the full shape of the data is a “U” nestled between Tg and Tm . Given the known difficulties in maintaining a constant, consistent temperature across the entire build area [24], understanding the sensitivity of a polymeric 0 material’s crystallization kinetics around the chosen temperature range is critical to affect desired performance properties and avoid out-of-plane warping. As shown in the figure, crystallization halftimes can vary over 100 minutes in the typical range for bed temperatures between 170-165 °C.91(1)(2)Interestingly, for a technology concerned with ensuring a slow crystallization to maximize particle coalescence and minimize out-of-plane warping, the “PBF-grade” materials have a shorter crystallization half-time than the control pellet at 160 °C.

The fit of Sintratec PA12 data to the Lauritzen-Hoffman equation indicates a short crystallization half-time even out to 175 °C. This is unsurprising given the known presence of filler material included for short-wave laser absorption that can act as a heterogenous nucleating agent. The industrial, white material DF2200 tracks closer to the control pellet. Based on Figure 3, the typical bed temperatures for printing nylon12, and potential build times of more than 12 h (720 min), one must conclude that it is unlikely for the entire part to “remain molten” for the entire build. Zhao, et al. report that, at minimum, the lower layers will begin to crystallize during PBF/LB based on their measured crystallization kinetics, which is reported as a deviation to what is commonly understood for the process [26]. Even considering that subsequent scanning of the laser re-heats and (at least partially) re-melts crystalline domains in a chosen layer of interest, Chatham, et al.

Even considering that subsequent scanning of the laser re-heats and (at least partially) re-melts crystalline domains in a chosen layer of interest, Chatham, et al. report that around 430 s (7.17 min) is the last time the temperature of a chosen layer of interest rises above the onset of melting for the laser parameter combination with highest energy tested [27]. Their work mapping cyclic thermal profiles measured during printing tensile specimens to crystallization kinetics indicates that useful molecular diffusion for coalescence is arrested due to crystallization after 500 - 600 s (8.34 - 10 min).Recall from the Introduction that the patent literature stated the magnitude of separation between observed crystallization and melting behavior in a dynamic DSC experiment was a good indicator of crystallization kinetics; slower crystallization kinetics would be accompanied by a larger separation between the two peaks. The presented data runs counter to this statement.

The presented data runs counter to this statement. The order of decreasing crystallization halftimes at 170 °C is (1) Sigma Aldrich control pellet, (2) DF2200, then (3) Sintratec PA12. Therefore, one would expect the same ranked order for decreasing degree92Figure 3: Time to reach 50% crystallized at indicated isothermal temperatures for DF2200, Sintratec PA12, and the Sigma Aldrich nylon-12 control pellet. Data were fit to both the original Lauritzen- Hoffman equation (dot-dash) [17] and an empirical version published by Patel and Spruiell (dashed line) [25].of melt/crystallization peak separation, but instead the Sigma Aldrich control pellet has the smallest extent of separation between peaks (See Figure 1). This underscores the importance of thorough consideration based in experimentation when considering new materials for PBF/LB.Physical GelationExperiments in the previous sections focus on the energetics of crystallization measured via DSC.

By contrast, physical gelation is measured via DMA as the mechanical response to crystallization. Molecular ordering into crystalline domains significantly increases a material’s modulus above Tg. While this process begins at local, discrete nucleation sites, subsequent growth of crystalline domains results in a gel-like network structure with the domains acting as physical crosslinks. The temperature at which this behavior begins can be determined via the frequency independent tan(δ); that is, the intersection of tan(δ) curves collected at multiple frequencies. The results from such an experiment are presented in Figure 4. The gelation temperature can be subsequently related to DSC experimentation at the same cooling rate to determine the extent of crystallization at the gelation temperature. When additionally combined with crystallization halftime data, candidate materials can be evaluated based on calculated time available for fusion, such as in Chatham, et al.

When additionally combined with crystallization halftime data, candidate materials can be evaluated based on calculated time available for fusion, such as in Chatham, et al. [27].Figure 4: tan(δ) collected at various frequencies plotted against temperature as the samples cool and crystallize from the melt state. The frequency independent junction of all curves is considered to be the physical gelation point.In addition to temperature for gelation, the retardation angle (δgel) is reported in Figure 4. This angle relates to the power law exponent n for typical mathematical representation of polymer complex modulus (G . The retardation angle is also related to the stiffness S of the gel.

This angle relates to the power law exponent n for typical mathematical representation of polymer complex modulus (G . The retardation angle is also related to the stiffness S of the gel. Correlations between gel stiffness and warping should be the subject of future work.) in the following manner:∗It should be noted that while the control PA12 pellet material from Sigma Aldrich presents a textbook example of frequency-independent tan(δ), the two samples of “PBF grade” powder show a wider range of crossover behavior smearing the gelation temperature. This is perhaps due to the complex nature of the fully formulated powder, which likely includes flow agents in addition to laser absorbers and crystallization controlling compounds. These additional components may prevent true frequency independence from occurring and may also introduce a deviation from power-law type behavior for the complex modulus.

However, the crossover spread remains in a tight temperature window of less than 3 °C, which is on the order of reliable temperature control for typical PBF systems.93Proposed New Processing WindowIn light of the discussed understanding of polymer crystallization, the authors propose the following revisions to the “PBF/LB-P Processing Window” from the common diagram based on vague patent language:Use Tmfor target minimum threshold temperature for laser scanning.Use Tgel obtained through DMA to determine the threshold for cooling and the critical percent crystallinity.Use t /and in situ Ttprofiles to estimate the time to reach critical percent crystallinity.These revisions are combined into Figure 5 for the DF2200 (white) and Sintratec PA12 (black) PBF/LB powders in addition to the Sigma Aldrich nylon-12 pellet control.12()The composite figure includes more and more relevant information summarizing polymer melting, crystallization, and degradation behavior than traditionally included when discussing polymers in the PBF/LB manufacturing context.

The overlaid data simply summarizes polymer behavior, but additional interpretation is needed to translate the results into the PBF/LB manufacturing context. Certain aspects are true physical thresholds, like the equilibrium melting temperature; however, many aspects are time-dependent and kinetic effects must be carefully considered when choosing process parameter settings.94Figure 5: Proposed new “PBF/LB-P Processing Window” diagrams for (a) Sintratec PA12 (black) powder, (b) DF2200 (white) powder, and (c) Sigma Aldrich nylon-12 pellet control. Diagram comprises DSC first heating and cooling, TGA, crystallization half-time, equilibrium melting temperature, and physical gelation temperature.95Note that the observed onset of melting is less than (or very close to) Tgel for all materials studied. Therefore, the Tb must be set below Tgel, which certainly allows for physical gelation to occur during printing.

Therefore, the Tb must be set below Tgel, which certainly allows for physical gelation to occur during printing. Given the goal of prolonging the coexistence of the recently-laser scanned melt with the solid-never-melted surrounding powder, the arrow indicating crystallization halftime at maximum bed temperature is a key metric to consider. Even though the Sintratec PA12 powder has the fastest crystallization rate at any given temperature of the three materials studied, its combination with the higher observed onset of melting enables successful printing. Similarly, the lower Tgel of the Sigma control pellet contributes to a prolonged time available for fusion; this is increasing total especially true if ensuring the temperature of the scanned region reaches above and Tgel is the scientifically more rigorous version of what was originally desired through the “SLS process window” looking at the observed onsets ∆ of melting and crystallization from the dynamic DSC experiment.

Framing polymer crystallization during PBF/LB around equilibrium melting and physical gelation is more consistent with established polymer science literature, which helps promote a correct understanding of polymer response to process stimuli truly enabling data-driven decision making when synthesizing novel materials or adapting and formulating existing materials for PBF/LB.T. In many ways, the difference between Tm0The presented example new composite figure for nylon-12 does not significantly differ from the typical diagram created based on the original patent language. This should be expected given that nylon-12 has been the dominant material studied and commercialized based on those patents. The authors expect material behavior for materials beyond nylon-12 to more significantly differ between the more rigorous “new window” diagram and typical window.

The authors expect material behavior for materials beyond nylon-12 to more significantly differ between the more rigorous “new window” diagram and typical window. It will, perhaps, explain the ability to print PEK HP3 and PPS despite not following the traditional separation of melting and crystallization window as reported in the literature [28, 29]. It should also be noted that the proposed new window diagram still does not include a direct representation of warping or the “50 µm out of plane threshold” suggested by the DTM patents [11, 12]. This should be the focus of future work, even though this behavior will likely show a heavy dependence on the geometry of the specific toolpath for the given layer than a true material property.ConclusionsInterdisciplinary teams are becoming more and more commonplace as society looks to tackle larger, more complex and interconnected issues. Such collaboration is only possible with a common vocabulary.

Such collaboration is only possible with a common vocabulary. This manuscript highlighted several terms pertaining to the melting and crystallization phenomena of polymers in the PBF/LB context that have historically imprecise meanings leading to misconceptions of the governing physics of this AM process. The authors hypothesize that this language and terminology barrier is hindering development, adaption, and commercialization of PBF/LB-P as the misunderstandings may result in research teams dismissing polymers outright for not having a “wide sintering window” instead of pressing in to determine suitable processing conditions.The authors summarize the key concepts of equilibrium melting temperature, isothermal crystallization, non-isothermal crystallization, and physical gelation as they have been well- studied in polymer science and briefly connect these topics to PBF/LB.

NDAMENTALS OF STEREOLITHOGRAPHYJuly 1992Dr. Paul F. JacobsDirector of Research & Development3D Systems, Inc. Valencia, California1961.INTRODUCTIONintroduction of From early 1988 until J.une 1992 over 300 SLA-1, SLA-250, SLA-190 and SLA-500 units have currently operating in 20 been These machines, the countries rapid prototyping systems now in use.Ithas only been a little over four years since thethe first StereoLithography system,the SLA-1.sold by 3D Systems.onfive continents,amount to about 90 percent of allthe field of rapid prototyping is still quite even among existing users, knowledge of the most basic relationships of this new technology is known.Notwithstanding these figures,young.The author has been surprised to note thatis at best uncertain.For newcomers,evenlessIt is therefore appropriate, on the occasion of the 1992 Sol i d Free...

Form Fabri cation Conference in Austi n us to develop those fundamental relationships which In the transition of StereoLithography from an art to a science, it is natural that we should attempt to develop a model of the process. Although the mathematics may seem formidable to some readers, the physical model is actually quite simple. While requiring only three key assumptions, we shall derive to nineteen fundamental relationships, important conclusions.forform the £oundation of this technology.sevenleadingphotopolymer The the involves in good theoretical predictions at least to first agreement with numerous experimental measurements, order. a good physical understanding of the basic phenomena. This description avoids extreme complexity and inevitable loss of generality that likely result from either analytical or numerical attempts at an would it was even more accurate model.

This description avoids extreme complexity and inevitable loss of generality that likely result from either analytical or numerical attempts at an would it was even more accurate model. While advanced studies felt the operational level.interaction of some very complex physics andactinic photons withreactivechemistry.

While advanced studies felt the operational level.interaction of some very complex physics andactinic photons withreactivechemistry. the model described herein areNonetheless,ofTheprimary benefit of this analysis is the development ofthecontinue,thatthe materialpresented is certainly appropriate at2.GAUSSIAN LASER SCANNINGConsiderat constant photopolymer.an actinic laser beam being scanned in a straightvelocity,free surface of the We shall make three key assumptions:overa vatofline, liquid1.The photopolymer exponential absorption.resin obeys the Beer-Lambertlaw of2.The laser irradiance distribution is Gaussian.3.The resin transitions from the liquid phase to the solid phase at the so-called "gel point".197Let us now define a coordinate system as shown in Figure 1. laser the increasing x. scan axis, with positive y defined by the right hand rule.

laser the increasing x. scan axis, with positive y defined by the right hand rule. directly under the centerline of the laser scan positive normal define resin. We coordinates of Q' are zero.Here in the direction of The y coordinate is laterally orthogonal to the laser Thus, y = ° the z coordinate extends downward into the resin, and is measured To consider some arbitrary point Q(x,y,z) within the be the projection of Q onto the resin surface.

zis scanned directly along the x axis,axis.Finally,to the x-y plane of the free resin surface, where z = the origin,0.Let Q'(O,y,O)shall arbitrarily select the origin such that both thexandFrom the Beer-Lambertlaw (Reference 1),H(x,y,z)= H(x,y,O) exp[- z/Dp ](1)where H(x,y,z) the Depth" depth of resin which will reduce the irradiance to lIe (about 37 %) the surface irradiance.is the irradiance at any arbitrary point, H(x,y,O)is and Dp is the "Penetration that ofsurface irradiance at any point x,y,O,ofthe resin at the laser wavelength.Dp is defined asFurther, (asGaussian approximation),if we assume that the laser irradiance distribution is reasonably good this is generally a seen in Figure 2,then from Reference 2,H(x,y,O)= H(r,O)= Ho exp[ -2 r'2.aI Wo ](2)where Wo is the radius of the Gaussian beam, defined per Reference 2 at the 1/e a point (i.e.

+y2.(11)thus,=exp [-2x / Wo ] * exp [-2y / Wo ]2.'2.2.2-(12)SUbstituting equation (12)into equation (10), moving the factor as it is also not a function of is the contribution from - ~ to 0 is exactly then the total integral must be twice theexp [-2y al Wo t x, and further recognizing that since the integral of equation (10) symmetric about x = 0 (viz. equal to that from 0 to _ value from 0 to infinity. Therefore,] outside the integral,)E(y,Z)2.

on the laser scan axis) and z exposure(18)it is evident that the exposure reachesmaximum value on the free resin surface). = value is given by the wonderfully simple expression;0(i.e.This maximum laser202Emax=( 2 /1(1)"2 Pa../ Wo Vs I(19)~------ is our first important result. (19) maximum actinic laser exposure is:EquationIt shows that the*Directly proportional to the laser power. Inversely proportional to the product of the beam radius and the scan velocity. The proportionality constant is simply ( 2 {11')" = 0.7979 .. , a pure number~involving no empirical quantities whatever.5.THE PARABOLIC CYLINDERFor photopolymer resins, when the exposure is less than a critical E > Ec, resin value, the if undergoes at least partial polymerization. However, resin is at the so-called "gel point", corresponding to the transition from the liquid phase to the solid phase.

However, resin is at the so-called "gel point", corresponding to the transition from the liquid phase to the solid phase. Hence, we may solve for the locus of points , which are just at the gel least partially point. solidified while all points outside this boundary will still be liquid. Clearly, cross-sectional resulting boundary will then define the shape of a single laser cured photopolymer "string".EcWhen,thetheresin remains liquid.E =Ec,z = z*y = y*,andthislocus will be atAll pointsinsidetheThus, setting y = y* after some algebra we find,andz = z* whenE = Ec in equation (18),2Z. exp [ 2y* / Wo+z* / Dp ]= (2 /'IT )'Ia { Pi.

building elements of StereoLithography are3showsscanning the surface of a vat of liquid photopolyrner, Thus, actually parabolic cylinders, often referred to as "strings".thefundamental6 .THE WORKING CURVEcured it string by the symbol Cd. simply stated, is evident that axis.

simply stated, is evident that axis. the maximum cure depth will occur directly under the laser scan y* = 0 Thus, setting z* = Cd when in equation (21), we finally obtain the fundamental "Working Curve" equation of StereoLithography,as discussed in Reference 5;We may now define the maximum cure depth of a single laserFrom either Figure 3, or equation (21),y* = o.z* = z(max) = Cd whenOr,Cd=DpIn [ Emax / Ec] I(22)~------~This is our third important result,and is absolutely fundamentalto an understanding of this technology.Equation (22)shows that:The maximum actinic laser exposure.cure depth should scale as the naturallogarithm oftheA semi-logarithmic plot of Cd vs In Emax should result in a straight line relationship.

. , :.~ . · . · · ·L.L..~I1Ir 0 ooo.. . . c"tr \I!IiI..II: . . . : . . :I III ~. . . . . . ..~o. ~o.. , . . . . .r o7.THE CURED LINEWIDTHReturning to equation (21) and Figure 3, it is also clear that the maximum cured linewidth, Lw, will occur at the resin surface, where the parabolic cylinder has its greatest width.

Therefore, setting y* = y(max) = Lw /2z* = 0, we obtain after some algebra,whenLw=.~ , 2 Wo { In [ Emax / Ec ] }V..(23)In [ Emax / Ec the basic "Cured Linewidth" equation,SUbstituting forfrom equation (22),we obtainLw=Wo ~ 2 Cd / DpI(24)This is our~---- important result.It shows that:The the calculating numerical values, of the laser spot, not the diameter.cured linewidth of a string is directly proportional laser spot size at the plane of the resin surface.to In remember that Wo is the radiusThe cured linewidth is also proportional to the square root resin penetration of also be depth.

resin, the laser laser spot size, and cure depth.Thus, sUbject to system servo limits,powerfaster the scan speed for agivenThe laser the ,cure depth.laserscan velocity isinversely proportional to the increasing the spot size decreases resin andspot size.Thus,laser scan velocity for a given laser power,The with an increase ln the ratio Cd / this slowly than shallow cure depths.laser scan v 7locity decreases in an exponential manner For a given resin, Dp. is the reason why increased cure depths draw much moreAgain, model is the same pure number,theconstant of proportionality( 2 / 11 )'/1.=predicted by 0.7979 . . . .thisEquation (27)is currently incorporated into 3D Systems software. on Experimentally measured scan velocity automatic typically within a few percent of the desired values.

. . .thisEquation (27)is currently incorporated into 3D Systems software. on Experimentally measured scan velocity automatic typically within a few percent of the desired values. the of approaching 0.15 mil orIt the SLA-190, cure depths, algorithm, Further, experimental about +/- 4 microns).is the basis for all automatic laser scan velocity calculationsand the SLA-500.the SLA-250,laserthegenerated usingareresidualerrorsarethethelimitstechnique (viz.standard deviations of +j-9.DRAWING TIME PER UNIT AREAIn StereoLithography,the great majority of the laser drawing time to solidify the regions of parts interior to their the time required to draw the is generally a small fraction of the time required to complete any let us and width Figure 5 shows this rectangle,is spent "hatching" borders. borders the hatching process. arbitrary cross section. However, consider w, perpendicular to the scan direction.

whose area is simply A = Lw .Exceptfor very tiny parts,For complete generality we could analyzeto simplify the calculations, in the scan direction,a rectangle of length L,straight parallel vectors with a hatch spacing, hs , as occurs with the advanced finite building techniques acceleration and deceleration effects at the ends of each hatch vector, the time required to draw a single string of length LNow,letusperform the hatching operation by drawingSTAR-WEAVE or....WEAVE TM andNeglectingis simply:2081-hs+I T,VsA--LFI(;UftE S.....1wt=L / Vs(28)bywithin integer round-off,the number of such vectors, N ,givenN=w / hs(29)Thus,to close approximation the laser drawing time,by the following:t d,is given=N t=( w / hs ) ( L/ Vs )=A / hs Vs( 30)Finally,the hatch Thus,in the advanced build methods WEAVET~and STAR-WEAVE™ Wo taken proportional to the laser spotspacinghs=k Wo( 31)where spacing (27) and (31)ka constant,generally near 2,so that the optimum hatch SUbstituting equationsis of the order of the beam diameter.into equation (30) we obtain after some algebra:ted / A='/&(1/k) ( tr / 2) { Ec / P L} exp [ Cd / Dp ](32)This is our sixth important result.

Equation (32) within the accuracy of the approximations we have made;indicates thatdirectly time per unit cross-sectional area The proportional to the resin critical exposure, Ec. Resins with higher values of Ec will draw more slowly.drawingThe the lasers will reduce draw time, the system servo limit.drawing time per unit area is inversely proportional laserpower, P~.As one might expect,higherprovided one has notto power reachedThe increased cure depth to penetration depth ratio. reS1n with draw much more slowly.

reS1n with draw much more slowly. accuracy, WEAVE TA utilize reduced cure depths relative to the Tri-Hatch method, more individual vectors are required.drawing time per unit area increases exponentially with For a given depths in addition to imFroved part STAR former they actually draw faster even though manya specific value of Dp,increased cureThus,the advanced techniques WEAVE T andbecause210is the observation that Finally, and especially significant, at least to first order, the laser drawing time per unit area is independent of the laser spot size. This result, which is probably not substantiated experimentally for all the approved stereoLithography resins.intuitively obvious,hasbeenThelatter point leads to the seventh importantSince order, system, a Dynamic zoom (viz. range) reduce provide productivity for those laser power levels which exceed the system servoresult.

range) reduce provide productivity for those laser power levels which exceed the system servoresult. time is independent of laser spot size to firstlaser drawing then dynamic optical unless one is always operating at the servo limit. the ability to vary the laser spot size over a wide would reliability, gain limit.zoom is of limited utility in suchdynamicincreasesystem cost,increase the above analysis shows,system complexity,and,asonly10.SUMMARYWehavederived some of the most fundamental theserelationshipsof experimentally Higher order effects such as resin "bleaching", non finite acceleration and deceleration intervals, somewhat. are solidStereoLithography. confirmed to varying levels of accuracy.

are solidStereoLithography. confirmed to varying levels of accuracy. optical self-focusing, Gaussian laser modes, and Nonetheless, predominantly valid, understanding of this technology.beenEach ofresults hasradiation scattering,system servolimitations will altertrends results discussed in thisthepaperthe mathematicaland should be considered fundamental to aREFERENCES1.Kauzmann, W. ," AnIntroduction to Quantum ChemistryAcademic Press, New York, 1957, pp. 578-579." ,2.Siegman, A. E. ,"AnIntroduction to Lasersand Masers",McGraw-Hill, New York, 1971, Chapter 8, pp. 312-314.3.Sneddon, I.N., "Special Functions of Mathematical Physics and Chemistry", Oliver and Boyd, London, 1961, pp. 13-14.4.Sherwood, G.E.F. and Taylor, A.E., "Calculus", Prentice-Hall, Inc., Englewood Cliffs, New Jersey, Third Edition, 1958, pp. 310.5.Richter, J., and Jacobs, P.

RTHER DEVELOPMENTS IN PROCESS MAPPING AND MODELLING INDIRECT METAL SELECTIVE LASER MELTINGC. Hauser, T.H.C. Childs and M. Badrossamay School Of Mechanical Engineering, University of Leeds, UKAbstractThis paper advances previous reported work on the mapping and modelling of single tracks and layers produced in powder beds of tool steel and stainless steel powders by a CO2 laser. For single tracks it reports on predicted and simulated track masses. It validates the simulations, including the use of absorption close to 1.0 when cylindrical tracks are formed. It also reports on melt pool temperature calculations and estimated melt pool dimensions which are used, in conjunction with bed physical properties, to explain why the single tracks form as either continuous with a crescent shape cross-section, continuous with an elliptical section, discontinuously irregularly broken, discontinuously balled or only partially melted as scan speeds and laser powers change.

It then extends its scope, experimentally, to consider effects of scan spacing on single layer formation.IntroductionMetal powders may be processed either indirectly or directly by Selective laser sintering (sls). In the indirect process the laser melts a polymer binder mixed with the metal powder to create a shaped but weak green part. Strengthening occurs from subsequent furnace sintering and infiltration [1]. In the direct process, the laser melts at least a fraction of the metal powder in its path. Traditionally, the powder used is often a specially blended alloy or a powder mixture containing low melting point metal binder particles which braze the non molten structural particles. Component densities are often low and still require strengthening by further sintering, infiltration or hot isostatic pressing [2,3,4]. More recently, a new technique is emerging in which alloy powders and alloy powder blends are totally melted.

More recently, a new technique is emerging in which alloy powders and alloy powder blends are totally melted. Its name is selective laser melting (slm), and its aim is to create a strong part that is usable without further processing other than perhaps surface finishing. However, processing is much more difficult by this route and is still the subject of major research studies. Control of melt pool flow is one major issue, more difficult for standard alloys with small melting temperature ranges [5] than for specially developed alloys or powder mixes with a wide range [6]. Frequently, fully molten metal will refreeze as unconnected or loosely connected balls rather than continuous tracks [7,8,9]. There are a number of approaches to avoiding this. Some authors propose as high a scan speed U as possible, subject only to keeping the energy density P/(U) within a certain range, where P is the laser power [7]. Others suggest using low P and U to avoid balling [8,10].

Others suggest using low P and U to avoid balling [8,10]. Most successful processing is with special alloys, with parts built up on a substrate, but that limits the areas to which the processing may be applied.The work of which this paper is part aims to process standard alloys without supports. This paper records the track forms seen in different (P,U) conditions when single tracks of ferrous alloy powders are melted by a continuous wave CO2 laser, scanning at low speeds. Depending on laser power and scan speed, tracks are formed either fully melted, partially melted or not melted at all. The fully melted tracks are either continuous or balled. An understanding of448what causes which form is then developed. A thermal model which enables the amount of melted powder to be predicted is tested on experiments and used to estimate the temperature field created in and around the melt pool.

A thermal model which enables the amount of melted powder to be predicted is tested on experiments and used to estimate the temperature field created in and around the melt pool. The model is then briefly applied to the simulation of raster scanning multi-track, single layer coupons, to explore the changes associated with melting powder in close proximity to melted and solidified powder.The Finite Element ModelThe finite element thermal, powder melting and densification model which is used for thermal and post processing calculations is discussed. The methods have been described before, both for laser sintering of polymers [11,12] and for recent developments in quantifying temperature fields created in metal slm [13,14]. It is the continual development of the latter which is the subject of review here.The calculation is a time stepping one. To start, melted tracks or layers may already exist in a bed. The laser is turned on and the first time step ∆t started.

To start, melted tracks or layers may already exist in a bed. The laser is turned on and the first time step ∆t started. A finite element calculation updates the bed temperature T(K) over the time interval ∆t. The bed density ρ (or porosity ε) is updated according to a (T,t) densification law. This leads to bed geometry change. The cycle is repeated as often as is required by the specified laser scan pattern. In the temperature calculation, thermal conductivity in the melted material and powder bed is assumed to vary with porosity as observed experimentally for copper [15]:k=ks) (( 1ε1−+a k)78.0εwhere ks is the solid conductivity and ak is chosen to fit the measured initial bed conductivity to its measured initial porosity.

ks and heat capacity Cp are assumed to vary linearly with temperature:ks=k, os+Ta , sk,Cp=C, op+Ta C.Evolution/absorption of latent heat L on melting/solidifying over the solidus/liquidus range TS to TL is treated by the temperature recovery method [16]. The move to metal track modelling mainly requires a shift from two to three dimensional modelling. However, some 2D results of metal powder bed melting simulations are reported in this paper. A five hundred-fold reduction in computation time has been achieved by changing the finite element solver from the previously used LUD method to a sparse matrix method [17]. Powder bed density is assumed to change from its initial value to that of solid material as temperature increases from TS to TL.

Powder bed density is assumed to change from its initial value to that of solid material as temperature increases from TS to TL. Based on viscosity dependence on solid fraction X in mushy state forming [18], it is supposed that there is no densification for X >> 0.5 but as X reduces below 0.5, densification increases instantaneously to a value depending on X. An empirical equation with these characteristics is, in terms of porosity:( εε1−powder)={ tanh( am[ 1−X] }n )am = 5 and n = 4 give significant densification for X < 0.7 and have been used here. X has been assumed to change linearly from 1 to 0 as T increases from TS to TL. The model does not follow convection of material in the melt pool. Densification is assumed to result in sinking of solidified material below the surface of the bed to form a crescent section track.

The model does not follow convection of material in the melt pool. Densification is assumed to result in sinking of solidified material below the surface of the bed to form a crescent section track. However, total cross- section area and hence mass m per unit length is calculated, as is the equivalent diameter D449(1)(2)(3)supposing the track is cylindrical. Thermal conductivity of and absorptivity α of laser energy into the powder bed are inputs to the thermal model. An approach to determining these, based on theory in [19], is used here. If heat αP∆t is instantaneously absorbed at a point on the surface of a powder bed, the maximum temperature rise at a depth z below the surface and the time taken for that to occur are:∆Tmax=147.0tP ∆α 3ρzC p,t∆Tmax=ρzC p 6 k2These equations are correct to better than 5% for laser beams of diameter 2a irradiating a surface for the finite time ∆t provided z/a > 4 and ∆t /t∆Tmax < 0.1.

A minimum estimate of absorptivity αmin is obtainable from the track melting experiments themselves. Equating heat absorbed to heat needed to melt the track gives:minα{( ( TCmUP = Lp)−T o)+}Lwhere To is ambient temperature. Consideration of α values is of major importance in this paper.ExperimentationMaterials. Three types of gas atomised powder have been used in this investigation: M2 and H13 standard tool steels and a stainless steel powder which is a high carbon content variant of type 314s that is used for high temperature super plastic forming moulds. The composition of each and their melting data is given in Table 1. All powders were obtained from Osprey Metals Ltd, UK and were supplied with a particle size distribution of -150 + 75µm and -75 + 38µm. No heat treatments, additives or fluxes or powder pre-heating was used in this work.Equipment. The slm equipment has been described before [13,14].

No heat treatments, additives or fluxes or powder pre-heating was used in this work.Equipment. The slm equipment has been described before [13,14]. A 120mm X 150mm tray, which contained the levelled powder bed to a depth of 7mm, was used for all single track and layer tests. The tray was placed within a custom built chamber able to be evacuated to 50mbar gauge pressure then back filled with a process gas to a pressure typically 30-50mbar over atmospheric to minimise leakage into the chamber. Before a series of tests the back filling cycle was repeated twice and the system was left to stabilise for 15mins. During stabilisation and testing, the gas flow rate through the chamber was 3 litres/min. In most of the tests reported here the process gas was 99.9% pure argon, but in one set of conditions melting was carried out in argon to which 10% hydrogen was added.Melting Experiments.

In most of the tests reported here the process gas was 99.9% pure argon, but in one set of conditions melting was carried out in argon to which 10% hydrogen was added.Melting Experiments. Single tracks and layers have been melted, by a scanning CO2 laser beam, in the surface of beds made from gas atomised powders of composition and size fraction reported above. Beam diameters at the powder bed surface were 0.6 mm and 1.1 mm. Delivered powers and scan speeds have ranged from 10 to 200 W and 0.5 to 50 mm/s for single tracks and 0.5 to 500mm/s for layers (see Table 2). For the layers the scan spacing’s were 0.15mm, 0.3mm and 0.45mm (25%, 50% and 75% of the beam diameter respectively). The tracks were nominally 30mm in length and 7mm apart and the single layers were 15mm x 15mm square. After a test, the tray was removed and the tracks and layers photographed in situ.

The tracks were nominally 30mm in length and 7mm apart and the single layers were 15mm x 15mm square. After a test, the tray was removed and the tracks and layers photographed in situ. Those tracks which were continuous and those layers which remained intact when handled were removed from the bed with tweezers, brushed to remove excess powder as best as possible, and weighed. The actual lengths, cross sections and widths and depths of single tracks were measure with callipers.450(4)(5)Simulations. The finite element model was used to simulate in the same conditions as the experiments. The main calculated quantities were the temperature distribution in the powder bed and the mass per unit length of melt tracks. The temperature distribution enabled length, width and depth of the melt pool to be estimated.

The temperature distribution enabled length, width and depth of the melt pool to be estimated. Two estimates of mass per unit length were calculated, an upper one for which the mass was taken to be of all that powder which experienced an increase in density of more than 100 kg/m3and a lower one which included all material that densified to within 300 kg/m3 of fully dense. The estimated masses were not sensitive to small changes in these chosen boundary values. The thermo-physical properties of 314S-HC, M2 and H13 in the solid state that were used in the simulations are listed in Table 3. Linear approximations to thermal conductivity and specific heat were obtained from data between about 100°C and 700°C in [20], though it was assumed that the specific heat of M2 was the same as that of H13. TS and TL for M2 and H13 were calculated using ThermoCalcTM software and the Scientific Group Thermodata Europe (SGTE) Solid Solution (SSOL) database.

TS and TL for M2 and H13 were calculated using ThermoCalcTM software and the Scientific Group Thermodata Europe (SGTE) Solid Solution (SSOL) database. Phase diagrams have been published previously [21]. TS and TL for 314S-HC were estimated from a conversion of its composition into Ni and Cr equivalents, according to formulae given in [22]. Values for L were obtained from the rule of mixtures and the latent heats of melting of the main alloy components Fe, Cr and Ni (again taking M2 and H13 to have the same values). All the materials were assumed to have a solid density of 7850 kg/m3.

Measurements of powder bed conductivity and absorptivity of laser radiation have been calculated from the theory based on equations 4 after heating beds with a 4mm diameter stationary 10W CO2 laser beam for 15s and measuring the variation of temperature with time 10mm below the beds surface, with a copper-constantan thermocouple. Beds were made in a similar manner to the main tests (to obtain similar density), but 30mm deep in a cylindrical container of 60mm internal diameter (the dimensions were tested to be effectively infinite from the point of view of heat conduction over the time period of the experiment). These experiments differed from the main tests in that they were carried out in laboratory air, rather than argon.Single Track Experimental and Simulated ResultsFive different qualitative forms of track were observed in the powder beds. Figure 1 illustrates these with examples formed at powers between 77W and 110W.

Figure 1 illustrates these with examples formed at powers between 77W and 110W. At the lowest scan speed, in an argon atmosphere, tracks (type A) were continuous and flat-topped or slightly concave in the surface of the powder bed. This is the shape expected from densification by melting, without material re-arrangement by convection or surface tension effects. With increasing speed, tracks (type B) remained continuous but became rounded and sank into the powder bed. This shape change, most likely caused by convection driven fluid motions (see later), causes a large denude area or trench to surround the track. In Figure 1, the rulers next to the type A and B tracks, with their 0.5mm divisions, give a scale to all of the tracks. Each ruler rests on the powder bed and therefore the region of dark shadow between the ruler and the track helps define track shape and gives an indication of the degree of sinking.

Each ruler rests on the powder bed and therefore the region of dark shadow between the ruler and the track helps define track shape and gives an indication of the degree of sinking. As the scan speed increases still further, the tracks (type C) became occasionally broken, although not always with the regularity of the example in Figure 1. Each fragment was still cylindrical in shape, and was often formed with a bobble of material at the start and, to a lesser degree, at the end of each melt fragment. The melt fragments were again surrounded by a denude area but did not sink to the same degree as track type B, with the characteristic bobble often rising above the level of the powder bed. At even higher speeds, the tracks became more regularly and frequently broken (type D). The termed ‘balled’ to describe the tracks became appropriate. Trenching, although to a much lesser degree, was still visible and the majority of the ball volume was sat up above the surface of the powder bed.

Trenching, although to a much lesser degree, was still visible and the majority of the ball volume was sat up above the surface of the powder bed. At the highest speeds, extremely fragile tracks were formed (type E). It will be shown later that these occur in conditions where maximum temperatures exceed TS but do not reach TL. They may therefore be described as partially melted.Process maps were created, to show the combinations of laser power and scan speed at which the different track types formed. Figure 2 shows four examples. In each, the dots mark the P,U combinations at which tests were carried out. The letters A to E show the regions of each track type. The region F is where laser power was insufficient to melt powder. All maps were qualitatively similar. Maps were also created for H13 and 314s -150/+38 powders using a 10% hydrogen addition in the process gas.

All maps were qualitatively similar. Maps were also created for H13 and 314s -150/+38 powders using a 10% hydrogen addition in the process gas. These maps were again similar and so were deemed unnecessary to be shown here.Figure 3, shows track cross section shapes to either side of the A-B boundary for H13 - 150/+75 powder. The change from flat to rounded tracks at crossing the boundary at constant power is accompanied unexpectedly by an increase in cross section area. This is discussed in detail in [14]. However, for the completeness of results it is important to re-emphasise the changes in calculated laser absorbtivity, αmin, (equation 5) from track mass/unit length data as the452A-B boundary is crossed. Values of Cp and L from Table 1 were used. Figure 4 presents the calculated values. The increase in track cross section shown in Figure 4 is associated with an effective increase in absorption. Values greater than 1.0, as shown for M2 are clearly impossible.

The increase in track cross section shown in Figure 4 is associated with an effective increase in absorption. Values greater than 1.0, as shown for M2 are clearly impossible. Perhaps inaccurate values of specific and latent heats have been used, or the tracks contain material that has not fully melted. Nevertheless, values are clearly greater than normally associated with absorption of C02 laser radiation by metal powders. These expected lower values are confirmed by the results of the further experiments using the stationary laser beam. The results of the further experiments are collected in Table 4. It also summarises the absorbtivity data from mass/unit length calculations, as well as powder bed densities and their thermal conductivities. Figure 4 also shows stills of tracks longitudinal section either side of the A-B boundary. The track nose is rounded.

Figure 4 also shows stills of tracks longitudinal section either side of the A-B boundary. The track nose is rounded. It has been discussed in [14] that that the shape of the leading edge of the track reflects the laser beam forward, creating a region of secondary melting. This was found to occur as speeds approx. greater than 2mm/s and may, in part, account for the increase in mass, and associated increase in absorbance, at the A-B boundary.Figure 1: Examples of five types of track A to E (all at nominally same magnification), from tests with -150/+75 µm M2 in argon atmosphere with a 1.1 mm diameter laser beam.Figures 5 to 8 report example main results of the simulations. Unless otherwise stated, material properties have been taken from Tables 3 and 4.

Unless otherwise stated, material properties have been taken from Tables 3 and 4. In A and E regions, α values determined from equation 4 have been used, while values from equation 5 have been used in the B, C, and D regions (except that for M2 and H13, α has been rounded to 1.0 rather than using the estimated 0.9 to 1.15 values). Values of ap (equation 4) have been chosen so that ∆zp is close to the mid-range size of powder particles in a bed (for example, for -150/ +75 µm powders, ap = 8 has been used, giving ∆zp ≈100 µm, depending on initial bed density).Figure 5 compares predicted and simulated track masses. It validates the simulation, including the use of α close to 1.0 when cylindrical tracks are formed. Here the simulations have taken the mass per unit length to include all material densified by more than 10 kg/m3.

It centres on a laser power range between 30W and 85W over which type E tracks are formed at a scan speed of 30mm/s. This range is translated from the map onto an adjacent graph which shows the calculated maximum temperature changes of the bed as a function of power, for a scan speed of 30mm/s and α values of 0.36 and 0.25 (0.36 is the value measured for M2 while 0.25 is the lower value measured for H13). The figure supports the hypothesis that the lower limit to region E represents conditions in which peak temperatures just reach TS while the upper limit corresponds to peak temperatures reaching TL.Figure 7 is concerned with conditions around the C-D boundary. Part (a) highlights two paths ab (crossing the C-D boundary at a constant power of 77W) and cd (along the C-D boundary) in the M2 process map. Part b records the calculated time in the melt and the length to diameter (L/D) ratio of the melt pool (assumed cylindrical).

Part b records the calculated time in the melt and the length to diameter (L/D) ratio of the melt pool (assumed cylindrical). Part (b) presents calculated maximum melt time for M2 material in a track and calculated length to diameter ratio of the melt pool, as scan speed is increased along path ab. Melt time reduces from close to 5s at the scan speed of 1 mm/s to 0.16s at the C-D boundary. Over that speed range, the melt pool length reduces from 4.8 mm to 2.4 mm (the melt time is the melt length divided by the scan speed) and the melt pool effective diameter (estimated from the track mass per unit length, on the assumption that the track cross-section is circular) reduces from 2.4 mm to 0.6 mm. As a result the length to diameter ratio increases along ab, reaching the value of 3.1 at the C-D boundary. Part (c) presents length to diameter ratios along the line cb, for M2 as well as the other two materials.

Part (c) presents length to diameter ratios along the line cb, for M2 as well as the other two materials. The results support the hypothesis that the C-D boundary is one of approximately constant melt length to diameter ratio. The constant lies in the range 2.8 to 3.3.Figure 8 is concerned with the A-B boundary. Part (a) reproduces a detail of the 314S-HC process map from Figure 2. Simulations have been carried out at constant power and increasing scan speed, along the three dashed paths (P = 77W, 110W and 143W) shown in the figure. Part (b) shows how the surface temperature gradient in the melt varied along these path, assuming the maximum temperature is at the centre of the melt pool (simulated) and the edge of the melt pool was at the temperature TL. The asterisks mark the conditions at the A-B boundary.

The asterisks mark the conditions at the A-B boundary. A-B may be a boundary of constant melt surface temperature gradient, of around 280 °C /mm to 320 °C/mm.455As the simulations do not model surface tension or convection driven fluid motions, and region A is one in which such motions do not occur, the measure is appropriate for conditions in region A, up to the A-B boundary.Figure 5: Experiment (•) and simulation (), P = 77WFigure 6: Process map for M2, -150/+75 µm powder size and 1.1 mm laser beam diameter, with temperature calculations at a scan speed of 30 mm/sFigure 7: (a) Part of M2 process map as Fig 10; (b) calculated tmelt and L/D along line ab; (c) L/D along line cd (also with values for 314S-HC and H13)456Figure 8: (a) A-B region process map for 314S-HC from Fig 6; (b) surface temperature gradient calculations along the dashed-lines of part (a)Single Layer ResultsThree different qualitative forms of track were observed in H13 powder beds (- 150/+75µm).

Figure 9 illustrates these with example layers formed at powers between 58W and 143W and for scan spacing’s of 0.15mm and 0.45mm. The processing gas was argon. At scan speeds below 150W (actual value dependant on laser power) layers (type 1) were fully melted with no to very little porosity. The porosity level was observed by holding the sample layer in front of a light box. With increasing speed, layers (type 2) were still fully melted but contained a much higher level of porosity (the actual level increased with increasing speed). Porosity levels were not measured but were judged to range between 60% and 95% of theoretical density.

Porosity levels were not measured but were judged to range between 60% and 95% of theoretical density. As the scan speed increased still further and as the laser power began to reduce, the layers (type 3) became very fragile and extremely porous, to a level where their removal from the powder bed became impossible without breakage.Process maps were created, to show the combinations of laser power and scan speed at which the different layer types formed for scan spacing’s of 0.15mm and 0.45mm (Figure 10a/b). Region 4 is where P/Us values were insufficient to melt the powder. A further two maps were also created (Figures 10c/d) which superimpose, for both scan spacing’s, layer types 2 then 3 with track types C then E respectively.

A further two maps were also created (Figures 10c/d) which superimpose, for both scan spacing’s, layer types 2 then 3 with track types C then E respectively. These groupings were chosen due to their close boundary proximity’s over the scanning conditions investigated, suggesting that perhaps porosity within a layer occurs due to the break up of the individual tracks.ConclusionsFive different track forms have been identified, depending on laser power and scan speed In region A, the tracks have a flat or slightly crescent shaped surface. The longitudinal track sections (Figure 4) suggest that the melt wets the surrounding powder. The melt ceases to wet the powder at the A-B boundary. This leads to a rise in laser power absorption, a growth of track section and the track sinking into the bed. In region C, tracks are occasionally broken. Even in B, breaks are sometimes seen, caused by interruption of the formation cycle.

In region C, tracks are occasionally broken. Even in B, breaks are sometimes seen, caused by interruption of the formation cycle. We suppose that in C, with its short melt times this happens more often. True instability starts at the C-D boundary. L/D ratios there are close to the value π expected of a Rayleigh instability. This further indicates that457the melt does not wet adjacent powder in regions B, C, D.

In region E, bounded by lines of constant P/U, partially melted track is evidence that maximum temperature lies between TS and TL.Layer type 1Layer type 2Layer type 3Direction of scanS = 0.15mmP(W) 143 110 58 U (mm/s) 10 80 150S = 0.45mmFigure 9: Examples of three types of layers 1 to 4 (all at nominally same magnification) from tests with -150/+75 µm H13 in argon atmosphere with a 0.55 mm diameter laser beam.200200)W(15012••••••••••• • • • • • • • • • • • • • •)W(150123••••••••••• • • • • • • • • • • • • • •r e w o P r e s a L100503••••••••••• • • • • • • • • • • • • • •••••••••••• • • • • • • • • • • • • • •4••••••••••• • • • • • • • • • • • • • •r e w o P r e s a L10050••••••••••• • • • • • • • • • • • • • •••••••••••• • • • • • • • • • • • • • •••••••••••• • • • • • • • • • • • • • •4••••••••••• • • • • • • • • • • • • • •0 100 200 300 400 5000 100 200 300 400 500Scan Speed (mm/s)Scan Speed (mm/s)(a)(b))W200Single Tracks (E)S = 0.45mm (3))W200S = 0.45mm (2)S = 0.15mm (2)(r e w o P r e s a L15010050••••••••••• • • • • • • • • • • • • • •••••••••••• • • • • • • • • • • • • • •S = 0.15mm (3)••••••••••• • • • • • • • • • • • • • •(r e w o P r e s a L15010050••••••••••• • • • • • • • • • • • • • •••••••••••• • • • • • • • • • • • • • •••••••••••• • • • • • • • • • • • • • •••••••••••• • • • • • • • • • • • • • •Single Tracks (C)••••••••••• • • • • • • • • • • • • • •0 100 200 300 400 500Scan Speed (mm/s)0 100 200 300 400 500 Scan Speed (mm/s)(c)(d)Figure 10: a/b H13 process maps for single layers with scan spacing’s of 0.15mm and 0.45mm respectively and c/d track types C and E superimposed onto layers types 2 and 3.458References[1] K.W.

rther Exploration of Multi-Material Fabrication Capabilities of Ultrasonic Consolidation Technique J.O. Obielodan, B.E. Stucker, Utah State University.AbstractThe increasing interest in engineering designs involving parts with multiple materials,and function specific members has placed more demand for technologies to fabricate such parts.This work discusses results of further exploration of multi-material freeform fabrication usingultrasonic consolidation. Various combinations of materials, including Titanium, Silver,Tantalum, Aluminum, Molybdenum, stainless steel, Nickel, Copper, and MetPreg have beenstudied. Some were found to be effective as a suitable intermediate layer between difficult to joinmaterials. Elemental Boron particles were added in situ between selected materials to modify thebonding characteristics.

Elemental Boron particles were added in situ between selected materials to modify thebonding characteristics. Microstructures of deposits were studied to evaluate bond qualities.Results show evidence of good bonds between various combinations of materials, thusillustrating increasing potential for multi-material freeform fabrication using ultrasonicconsolidation.1.IntroductionUltrasonic consolidation is a solid-state fabrication process that combines ultrasonicmetal welding and layered manufacturing techniques to produce three-dimensional freeformobjects. The process uses the power of high frequency ultrasonic vibration at low amplitudes tobond thin foils of materials to form solid objects. It combines normal and oscillating shear forceson mating foils and the resulting friction forces between the materials to fracture and displacesurface oxides from the materials.

It combines normal and oscillating shear forceson mating foils and the resulting friction forces between the materials to fracture and displacesurface oxides from the materials. The exposed atomically clean surfaces are then brought intodirect contact under modest pressure and temperatures that are less than half of the meltingpoints of the materials. The materials are thus metallurgically bonded [1]. Fractured oxides andsurface impurities in the materials are distributed in the bond zone. The process combines thelayer-by-layer addition of foils with contour milling using the integrated 3-axis CNC machiningfacilities to produce desired component geometry. It is therefore both an additive and subtractiveprocess. Fabrication involves the generic freeform fabrication additive process chain in which asolid CAD model is numerically sliced and the thin layers sequentially sent to the fabricatingmachine to build the part from bottom up.

Apart from removing the substrate upon which thedeposition is made after fabrication is completed, no further machining of the part is required,making it a near-net shape fabrication. If the substrate is part of the final component, it is netshape fabrication.Some notable advantages of the solid state UC process are as follows [1].No high-temperature process-associated safety hazards.No atmospheric control is required.As low temperature is involved and the volume of material affected is small, lessenergy is needed.Embrittlement, residual stress, distortion and dimensional changes are greatlyreduced with the low temperature processing.The UC machine consists of a welding horn, also known as a sonotrode, that exerts anormal force and the oscillatory high-frequency vibration on the materials to be welded. Weldingtakes plate on a substrate fixed on a heated plate.

Weldingtakes plate on a substrate fixed on a heated plate. The UC machine is designed for automatic foilmaterial feeding, but materials can also be fed manually. Figure 1 shows the schematic view ofthe ultrasonic consolidation process. The primary process parameters are vibration amplitude,temperature, welding speed, and normal force [2]. Other parameters that can affect weld qualitiesinclude sonotrode roughness, material surface finish [3], and side-by-side foil positioningaccuracy with respect to the automated material feed system [4].Figure 1: Schematic of UC processUltrasonic consolidation is applicable for rapid tooling for injection molding, extrusion,vacuum forming tools and others. It is also been used for fabricating tools with conformalcooling channels [1].

It is also been used for fabricating tools with conformalcooling channels [1]. Previous work have demonstrated other potential applications of UC,which include honeycomb structures [5], embedding shape memory alloy (SMA) fibers andsilicon carbide fibers in aluminum matrices [6-9], and embedded electronics [10]. While theprocess has been widely used for single material fabrication with aluminum alloys, only a fewresearchers have demonstrated its capabilities for multiple material fabrications. The multi-materials capabilities of UC was demonstrated by Janaki Ram et al. [11] in their work in whichcopper, brass, nickel, inconel 600, AISI 347 stainless steel, stainless steel AISI 304 wire mesh,MetPreg, and aluminum alloy 2024 were individually welded to aluminum 3003 H18 materials.Domack et al., also in their work [12], demonstrated the capability of UC for graded materialscomposition fabrications using titanium and nickel alloys.

In this present work, the capabilitiesof UC to fabricate multi-material structures are further explored. Suitable combinations ofMolybdenum, Tantalum, Nickel, Stainless Steel 316L, Silver, MetPreg, Copper, Aluminum1100-O, Aluminum 3003 H18, and Aluminum 6061-O, were bonded using aluminum alloy3003-H14 and Aluminum alloy 6061-T6 substrates. Boron powder was added in situ for some ofthe material combinations.Engineers and designers desire to harness the benefits of combining a variety of function-specific materials where they are needed, and the geometrical complexities offered by the UCprocess to fabricate these structures.

The applications of multi-material functional structures arediverse, including surface protection with corrosion or wear resistant materials, radiationshielding, and combining electrical insulators with highly conductive materials for use inaerospace, automobile, ship building, nuclear, electronics, industrial machineries and otherindustries.2.Experimental WorkA Solidica FormationTM machine was used for the experimental work. Although themachine has capability for automatic foil feeding, all foils used were manually fed exceptaluminum 3003 foils. Materials of 40x20mm size of variable height, depending on the number offoils and their thicknesses, were deposited on the Al 3003 and Al 6061 substrate materials. Thetwo substrates were of 355 x 355 x 12 mm size. For each deposit, several layers of materialswere welded to demonstrate their weldability within the current limits of the primary weldingparameters of the UC machine.

For each deposit, several layers of materialswere welded to demonstrate their weldability within the current limits of the primary weldingparameters of the UC machine. Different arrangements of foil stacking for the materials usedwere experimented with. In each material combination, the welding parameters used for eachlayer was dependent on the material to be welded at any instant, so for each material, the mostsuitable welding parameters were used.

The optimum process parameters for Aluminum alloy3003 were obtained in previous work by Kong et al and Janaki Ram et al [2-3]. While work isstill ongoing to determine the optimum parameters for most of the other materials in differentcombinations, the parameters used in the present work were found to work well for therespective materials in Table 2. Aluminum alloy 1100 was generally used as an interlayermaterial between difficult to weld materials. The interlayer was manually placed between thedifficult to join materials and the sonotrode run on the topmost material to weld them together.The Al 1100-O interlayer material was found to bond well with most of the materials used in thisstudy. In cases where boron powder was added at the interface of two materials, the powder wasthoroughly mixed in water and a brush was used to apply the mixture onto the surface of thesubstrate or already welded foil.

After moisture evaporation, the foil to be welded on it ismanually placed for welding. Mixing the boron powder in water was found to make it adheremore effectively to the substrate before welding than applying loose, dry powder.Small samples of the deposited materials were mounted and polished according tostandard metallographic procedures and observed under an optical microscope. The bondqualities between the foils of different materials were qualitatively evaluated.3.ResultsMicrographs of bonded materials are shown in Figures 2 to 8. The description of thewelded foils in each figure is such that, the material that is welded directly on the substrate is thefirst, followed by the next material, and continuing to the topmost foil material.

[11] highlighted the dominant factors influencing good bonding between two foils ofmaterials. The ability to plastically deform the foils under the action of the normal and oscillatingshear forces acting at the interface of the mating foils is of paramount importance as it helps inbreaking the hard surface oxides and repeatedly deforming the surface asperities therebyexposing atomically clean surface for metallurgical bonding between the mating foils during theweld cycle. Successful welding between two mating foils can be a measure of how well thesurface oxides of the foil materials can be removed as well as the ease of surface deformation.From the results presented, it can be observed that the relatively soft materials generally bondedwell to each other. This is because, their surface oxides were easily broken up and displacedalong with the fact that they were relatively easier to deform under the influence of the operatingforces.

This is because, their surface oxides were easily broken up and displacedalong with the fact that they were relatively easier to deform under the influence of the operatingforces. It is also worthy of note that most of these softer materials have face centered cubic(FCC) crystal structures while the harder ones have body centered cubic (BCC) and hexagonalclose pack (HCP) structures at UC processing temperatures. So far limited success has beenachieved in bonding softer materials to hard ones. All the aluminum alloy series used in thisexperiment bonded well with all the hard materials tried, whereas silver, copper and nickel didnot bond well with molybdenum and tantalum directly. Thus, in addition to welding parameters,material composition and combinations play an important role in determining good bonding.Attempts to bond molybdenum and tantalum, representing hard materials to themselves both assimilar and dissimilar materials, have been unsuccessful.

Observations from welded materialcombinations are discussed in the following subsections.4.1 Silver/Copper/Nickel weldingIn this combination of materials, two silver foils were successfully welded to each otheron the Al3003 base plate as shown in Fig.2. Copper foil also bonded well to both silver at thebottom and nickel at the top. From the micrograph, it can be seen that the silver to silver weldhas a very good linear weld density (a measure of bond quality [3]), as there are no visibleinterfacial defects between the two layers. Silver welded well at 150oF as well as at 300oF,although it underwent a high rate of oxidation at 300oF. Except when bonding with materials thatwill significantly conceal the surface from atmospheric oxygen, it is better to weld Silver at150oF. Copper to silver and copper to nickel foils also bonded very well to each other, as can beseen in Figs. 2a and 2b.

Copper to silver and copper to nickel foils also bonded very well to each other, as can beseen in Figs. 2a and 2b. From the micrographs, good welds were obtained between copper andnickel both at the top and bottom. In Fig.2c, nickel, copper and silver are shown individuallysandwiched between aluminum alloy 6061-O foils to demonstrate their good weldability to thisaerospace grade material using UC. Al 6061-O foil bonded well with the Al 6061-T6 substratematerial with no visible defects. All the materials welded in Fig. 2 show good bond qualities.4.2 Molybdenum/Al 3003-H14/Copper WeldingFigure 3 shows the micrographs of moderately bonded combinations. Molybdenumwelded to aluminum alloy 3003 H18 is shown in Fig.3a, copper with aluminum alloy 1100-O asan interlayer material in Fig.3b and to aluminum alloy 6061-O in Fig.3c. Molybdenum is a hard,wear and corrosion resistant refractory metal with high temperature strength, and resistance toplastic deformation.

Molybdenum is a hard,wear and corrosion resistant refractory metal with high temperature strength, and resistance toplastic deformation. As can be seen in Fig.3a, molybdenum bonded well with Al 3003 witheither material at the top or bottom. It is also bonded well to copper when using an Al 1100interlayer material. Within the limits of the welding parameters of the UC machine used,molybdenum could not be successfully joined to copper directly. With the interlayer materialhowever, very good joining was achieved. Molybdenum-copper laminated materials haveapplications in thermal management for electronics packaging [13].

With the interlayer materialhowever, very good joining was achieved. Molybdenum-copper laminated materials haveapplications in thermal management for electronics packaging [13]. The bond betweenmolybdenum and Al 6061-O shown in Fig.3c is not as good as those in Figures 3a and 3b.However, with further welding parameter optimization, there may be a possibility of achievinggood bonding between these two materials.4.3Tantalum/Aluminum WeldingMicrographs of tantalum foils ultrasonically welded with aluminum alloys 3003-H18 and6061-O are shown in Fig. 4. The tantalum foils were in the as-rolled and tempered form. Themicrographs show very good linear weld densities between tantalum and the two aluminummaterials. The sandwiched tantalum between the two aluminum alloys show that good bonds areachieved with either material at the top or bottom of the welds. Tantalum is a refractory metalwith good wear and corrosion resistance.

Tantalum is a refractory metalwith good wear and corrosion resistance. It can be ultrasonically welded to a material for surfaceprotection against wear and corrosive environmental conditions. Tantalum is also used forradiation shielding in nuclear applications [14].4.4 MetPreg/Copper WeldingMetPreg was fully bonded with copper on an aluminum 3003 substrate, as shown in themicrographs in Fig. 5. The deposit is a combination of an aluminum metal matrix material withvery hard reinforcing alumina fibers, joined to copper, a good electrical and heat conductingmaterial. The dual material deposit can be used in applications requiring good heat or electricalconductivity underlying materials with a hard, wear resistant surface.

The dual material deposit can be used in applications requiring good heat or electricalconductivity underlying materials with a hard, wear resistant surface. The micro-hardness of theas fabricated surface of the MetPreg on copper is 600Hv.4.5Titanium/Aluminum WeldingTitanium was successfully joined to aluminum alloys 3003-H18 and 6061-O byultrasonic consolidation as shown in the micrographs in Fig. 6. The micrographs show goodbond qualities between titanium foils and the aluminum alloys with either material at the top orbottom position. Titanium and aluminum have a wide range of applications in the aerospaceindustry.

Titanium and aluminum have a wide range of applications in the aerospaceindustry. As such; ultrasonic consolidation provides a unique fabrication technique for their dualmaterial freeform fabrication for functional structures in the aerospace industry, especially withaerospace grade Al 6061 materials.4.6 Titanium/Aluminum Welding with Embedded Boron ParticlesTitanium and aluminum alloy 3003-H18 welded well with embedded boron powder at theinterface as shown in the micrographs in Fig.7. The boron powder used has particle size less than5µm diameter. Plastic flow of aluminum and titanium foils around the boron particles is crucialto obtaining good bond. During welding, the oscillating motion of the vibrating sonotroderedistribute the particles at the interface of the welded foils, as such, uniform particle distributionwill be difficult to achieve. Embedded powder can be used to achieve local property controlwithin a structure.

Embedded powder can be used to achieve local property controlwithin a structure. It can also be used to fabricate particle reinforced composite materials,especially in cases where the UC particle embedment is an initial fabrication step before postdeposition heat treatments. The deposition shown in Fig.7 was subjected to post processannealing at 480oC for two hours and oven cooled. The result of Energy Dispersive X-ray (EDX)spot analysis, shown in Table 3 below at a 1µm point distance on the aluminum side from theboron powder, reveals that significant boron diffusion into the aluminum matrix took place.4.7 Nickel/Stainless Steel 314L WeldingFigure 8 shows a good bond between stainless steel 316L and nickel.

The austeniticstainless steel material with an FCC crystal structure demonstrates good weldability with nickel.The dual materials have applications in areas where strength and corrosion resistance is a majorrequirement.Table 3: Result of EDX spot analysis of a point in the aluminum side of the matrix5Concluding RemarksThe multi-material capabilities of ultrasonic consolidation have beenfurtherdemonstrated in this work. All FCC crystal structure materials used welded well with each other.Among the materials used, only aluminum alloys 1100, 3003, and 6061 welded well withmolybdenum, tantalum and titanium; the only non-FCC crystal structured materials. With furtheroptimization of the welding parameters for most of the material combinations, functional multi-material structures can be fabricated by ultrasonic consolidation.Acknowledgementsthe National Science of Naval Research (under Grant No. N000140710633) for financially supporting this work.

rther study of the electropolishing of Ti6Al4V parts made via electron beam melting Li Yang, Austin Lassell, Gustavo Perez Vilhena PaivaDepartment of Industrial EngineeringUniversity of Louisville, Louisville, KY 40292Abstract In this study, the effect of various parameters including the voltage, current, polishing time, temperature and electrode spacing on the electropolishing quality of Ti6Al4V samples made via electron beam melting was investigated using specialty designed research setups. The relationships between these process parameters and the surface roughness of the parts were established experimentally.Introduction In the previous study, the authors investigated the use of non-aqueous electropolishing as a potential method for the surface treatment of Ti6Al4V parts fabricated via electron beam melting (EBM) [1]. A surface finish of 6μm was achieved via experiments, although the uniformity of the surface polishing was rather unsatisfactory.

A surface finish of 6μm was achieved via experiments, although the uniformity of the surface polishing was rather unsatisfactory. Furthermore, through the investigation of relationships between several process control variables such as electrode spacing, polishing time and the surface finish, it was concluded that this method could potentially achieve efficient surface finish improvement for EBM Ti6Al4V parts and structures [1].Various literatures have reported the high sensitivity of the electropolishing process on various process parameters such as temperature, agitation and electrical excitation energy [2-5], which is largely due to the underlying polishing mechanisms. For a surface feature that possesses both macroscopic waviness (>1μm) and microscopic roughness (<1μm), which are driven by different mechanisms [3]. While macro-smoothing are primarily dominated by the electrical current level, the micro-smoothing is dominated by mass transportation of dissolved anions (i.e.

While macro-smoothing are primarily dominated by the electrical current level, the micro-smoothing is dominated by mass transportation of dissolved anions (i.e. diffusion). For most electropolishing practices, both macro- and micro-smoothing are desired, as macro- smoothing removes large surface roughness, and micro-smoothing result in brightening of the surfaces.

For most electropolishing practices, both macro- and micro-smoothing are desired, as macro- smoothing removes large surface roughness, and micro-smoothing result in brightening of the surfaces. Therefore, as diffusion is strongly influenced by both temperature and the fluid flow that carries the reaction species away from the electrodes, the electropolishing effects are affected consequently.Multiple works have investigated the electropolishing of titanium and titanium alloys using various electrolyte recipes such as methanol-sulfuric acid [4], methanol-buthanol-perchloric acid [6], hydrofluoric acid-lactic acid-sulfuric acid [7], perchloric acid-acetic acid [8, 9], perchloric acid-ethylene glycol monobutylether-methanol [10], sulfuric acid-hydrofluoric acid-acetic acid, sulfuric-phosphoric acid-hypophosphite-tartrate, anhydrous sulfamic acid-formamide and sulfuric acid-KF-ammonium bifluoride [11].

During the electropolishing process, the polishing reaction and oxidation reaction as shown in Eq.(1) often take place simultaneously. While thepolishing reaction dissolves titanium atoms from anode parts, the oxidation (passivation) reaction forms a thin and uniform TiO2 layer on the part surface that limits the titanium anion diffusion into the electrolyte solution. As shown in Fig.1, for titanium the voltage-current electropolishing characteristic often exhibits an initial slope followed by a plateau stage in which the electropolishing current remains largely constant. It could be expected that the dominant polishing mechanism is macro-smoothing during the initial stage, which changes into a combined macro-/micro-smoothing mechanism that is desired for the process.

It could be expected that the dominant polishing mechanism is macro-smoothing during the initial stage, which changes into a combined macro-/micro-smoothing mechanism that is desired for the process. On the other hand, as shown in Fig.1, there exist a complex balance between various reactions during the plateau stage, including the diffusion of the dissolved titanium anion into the electrolyte through the passive oxide surface film under electrical potential, the build-up of the titanium oxide layer and the breakdown of this oxide layer.

Due to the high propensity of passivation, titanium alloys often requires highly corrosive acids such as hydrofluoric acid and perchloric acid in order to effectively break down the oxide layer during the polishing process, which pose serious health and safety hazards.Polishing: Ti(cid:1)Ti4++4e(1.1)Oxidation: Ti+2H2O(cid:1)TiO2+4H+4e-(1.2)Ti+2OH-(cid:1)TiO2+H2O+2e-(1.3)Fig.1 Typical electropolishing characteristic of titaniumThe use of non-aqueous electrolyte was suggested to provide a possible solution, as the existence of water strongly promotes the formation of stable titanium oxide layer [4]. Several anhydrous titanium electrolyte recipes have been developed, including the previously mentioned sulfamic acid-formamide [11] and ethanol-isopropyl alcohol-zinc chloride-aluminum chloride [12]. The latter was of specific interest to the authors as all the ingredients are commonly available and pose less overall health and safety hazards.

The latter was of specific interest to the authors as all the ingredients are commonly available and pose less overall health and safety hazards. It was reported in the previous study that this electrolyte does not eliminate the formation of oxide layers. Instead, a viscous surface layer consists of both titanium oxide and titanium salt was formed, which can be relatively easily removed via electrolyte agitation [12].In this study, an improved electropolishing device was designed and built, which allows for improved control with the temperature and the electrolyte agitation. Utilizing the improved device, experiments were performed to characterize the electropolishing process and to optimize the polishing parameters.Experimental setup The improved experimental setup is shown in Fig.2. The main reaction chamber was a custom built 70mmx85mmx85mm glass tank with two fluid inlet/outlet, which was fixed on a wooden substrate during the experiments.

The main reaction chamber was a custom built 70mmx85mmx85mm glass tank with two fluid inlet/outlet, which was fixed on a wooden substrate during the experiments. The electrodes were connected to an Instek PSW 160-14.4 multi-range DC power supply. The electrolyte agitation is driven by a Masterflex 7523-40 L/S peristaltic pump and circulates through a custom built Liebig condenser as the heat exchanger for heating/cooling of the electrolyte. The temperature of the water used for heating/cooling was controlled by a water tank with automatic heating/cooling control (not shown in Fig.2). The heating/cooling was designed to be indirect as it was found to be difficult to find an economical direct cooling solution with the high-purity ingredients used for the study.

The heating/cooling was designed to be indirect as it was found to be difficult to find an economical direct cooling solution with the high-purity ingredients used for the study. In addition, a thermocouple was placed into the electrolyte solution with glass protection in order to monitor the temperature variation during the electropolishing process.Fig.2 Electropolishing experimental setupDuring the experiments, the peristaltic pump was set at a constant flow rate of approximately 1.3L/min, which was considered sufficient as the electrolyte solution volume used for each experiment was approximately 400mL. As the heated water tank possesses a rather large heat sink, the temperature setup was done prior to the experiments to ensure sufficient heating of the electrolytes.

As the heated water tank possesses a rather large heat sink, the temperature setup was done prior to the experiments to ensure sufficient heating of the electrolytes. However, it was found during preliminary experiment that there still exist considerable temperature drift in the electrolyte solution even with the use of the cooling mechanism shown in Fig.2, which was largely due to the large voltage and current used in electropolishing. Therefore, the electropolishing process was set to pause for 60s for each 60s of polishing to allow for additional time for cooling, which was shown to be able to limit the temperature drift to within 10°C for all the experiments performed. The control of the on/off of the power supply was realized via a LabView program.For the electropolishing, Ti6Al4V bars with 20mmx8mmx2mm size were fabricated in Arcam S400 system using standard Ti6Al4V solid theme. All the samples were fabricated in the same orientation as shown in Fig.3a.

All the samples were fabricated in the same orientation as shown in Fig.3a. A grip end was also designed for each sample to facilitate the gripping of the parts during electropolishing. A copper foil was used as the cathode, with dimensions sufficiently larger than the processed samples. After polishing, the surface finish of the samples were measured by a Dektak 8 profilometer. All the measurements were done with a 2.5mm measurement length around the central area of the polished surface as illustrated inFig.3b. Beside the surface finish measurement, the weights of the samples were also measured before and after the electropolishing in order to evaluate the material removal.a. Sample designb. Surface roughness measurementFig.3 Electropolishing samples fabricated by EBMExperiments and Discussions The electropolishing characteristics of the experimental setup was evaluated at multiple levels of temperatures in order to determine reasonable ranges of work envelope.

Fig.4 shows the voltage-current curves of the experimental setup under three different temperatures ranging from 25°C to 53°C. It could be clearly observed that temperature has a significant effect on the electropolishing process. With only 30°C difference, the polishing speed could vary by 100-300% depending on the voltage level, which shows the importance for temperature control during this process. It was also found that very little reaction took place when the electrolyte temperature was set at 4°C, which can be readily explained by considering the fact that the Nernst potential is significantly influenced by temperature. In addition, no plateau stage was observed, which agreed with results from previous study [12]. This has two potential explanations. First, due to the lack of formation of the stable oxide surface layer, the process was not hindered significantly. Second, the reaction might had not reached the diffusion rate limit due to the sufficient electrolyte agitation.

Second, the reaction might had not reached the diffusion rate limit due to the sufficient electrolyte agitation. After referring to the previous results, it was determined that further electropolishing experiments would be performed at the 60-80VDC range.Fig.4 Electropolishing characteristic curve for the experimental setup In the evaluation of effects of different process parameters on the electropolishing effects, the experiment design is shown in Table 1. For all the experiments, the electrode spacing (i.e. distance between the cathode and the anode) was kept constant at 5mm, which was identical to the setup used for process characterization.

For all the experiments, the electrode spacing (i.e. distance between the cathode and the anode) was kept constant at 5mm, which was identical to the setup used for process characterization. Two levels of voltage (60V, 80V), two levels of temperature (27°C, 38°C) and three levels of polishing time (5min, 10min, 20min) were investigated, and a full factorial design was performed.Sample 1 2 3 4 5 6 7 8 9 10 11 12Voltage (V) 60 60 60 80 80 80 60 60 60 80 80 80Temperature (C) 38 38 38 38 38 38 27 27 27 27 27 27Time (min) 20 10 5 20 10 5 20 10 5 20 10 5Table 1 Electropolishing experimental designFig.5 shows the samples after the electropolishing. The effect of electropolishing was visually discernable. With longer polishing time, the loss of shape also becomes more significant.

The effect of electropolishing was visually discernable. With longer polishing time, the loss of shape also becomes more significant. Temperature also appeared to be clearly more dominating compared to voltage.Fig.5 Samples after electropolishingFig.6 shows the surface finish measurement results for the polished surface under each process conditions. It could be seen that when the sample was process at 80V+38°C condition, a minimum surface roughness of 2.3μmRa was achieved, although it did not appear with the 20min sample. In fact, it was observed in most samples that the polishing time did not appear to have a monotonous effect with the surface finish. This might be caused by the repeated use of the electrolyte solution for each voltage and temperature combination, as well as chemical pitting [3] in the case of the 80V+38°C group.

This might be caused by the repeated use of the electrolyte solution for each voltage and temperature combination, as well as chemical pitting [3] in the case of the 80V+38°C group. The repeated use of solutions due to the limitation of available resources would result in gradual change of the solution composition in the electrolytes, which might potentially result in the promotion of stable oxide layer formationat specific compositional ranges. Although preliminary study by the research group did not observe significant change with the voltage-current electropolishing characteristics between fresh and used electrolyte solutions, further studies are needed to clarify about this. Another possible cause might be attributed to the measurement process. As shown in Fig.7, the surface profile measured by the profilometer for the 80V+38°C+20min sample actually exhibit very smooth topology.

As shown in Fig.7, the surface profile measured by the profilometer for the 80V+38°C+20min sample actually exhibit very smooth topology. However, due to the loss of geometry, the sample could not be leveled well enough during the measurement, which result in a macroscopic profile skew that could become significant over a 2.5mm measurement range.Fig.6 Surface finish of processed surfacea. As-received sampleb. Polished sampleFig.7 macroscopic skew of the sample surface finish measurementIt was also interesting that the back surface of the samples were also polished. Fig.8 shows the surface roughness of the back surface of the samples. The surface finish on the back surface exhibited more consistent trends in relation to the polishing time, and the best surface finish of 0.6μmRa was obtained from the 80V+38°C+20min sample.

The surface finish on the back surface exhibited more consistent trends in relation to the polishing time, and the best surface finish of 0.6μmRa was obtained from the 80V+38°C+20min sample. Also, no significant pitting was observed, which suggests that the reaction on the back surface might be more controllable.Fig.8 Surface roughness of the back surfaceThe measurement results of weight loss for all the samples is shown in Fig.9. The results of weight loss verified that the material removal effect is proportional to the polishing time, which suggests that chemical pitting might in fact occurred in most of the samples. The weight loss for the high voltage and high temperature group was very significant, with a maximum loss ratio of about 30%. On the other hand, it appears that the material removal rate could be accurately predicted.

On the other hand, it appears that the material removal rate could be accurately predicted. Therefore, it might be possible to accommodate this during the design of the samples in the future.Fig.9 Weight loss of electropolishingConclusions In this study an improved electropolishing device was implemented, which realized better control of electrolyte agitation and temperature for the electropolishing process. After implementing the new setup as well as control process pause, a 10°C temperature drift was observed during each electropolishing experiment. Through voltage-current characterization it was speculated that the electropolishing performed in this study was dominated by macro- smoothing mechanism. The electropolishing was able to achieve significant surface finish improvement from over 20μm with original parts down to <1μm, although the uniformity of the polishing is still low.

The electropolishing was able to achieve significant surface finish improvement from over 20μm with original parts down to <1μm, although the uniformity of the polishing is still low. Further studies are needed to optimize the process and to demonstrate themechanical benefits of the processed samples via fatigue testing, which is currently undergoing by the authors.Acknowledgement The authors would like to acknowledge the support from Rapid Prototyping Center (RPC) at University of Louisville, as well as various personnel including Dr. Gamini Sumanesekera, Dr. Thad Druffel and Dr. Mahendra Sunkara. This work is funded by the Intramural Research Incentive Grant (IRIG) from University of Louisville.Reference [1] L. Yang, H. Gu, A. Lassell. Surface treatment of Ti6al4V parts made by powder bed fusion additive manufacturing processes using electropolishing. Proceedings of Solid Freeform Fabrication (SFF) Symposium, Austin, TX, 2014.[2] H.-J. Liu, Q. Xu, C.-W. Yan, Y-.Z. Cao, Y.-L.

RTHER UNDESTANDING OF Ti-6Al-4V SELECTIVE LASER MELTING USING TEXTURE ANALYSISM.Simonelli*, Y.Y. Tse*, C. Tuck†Department of Materials, Loughborough University, LE11 3TU, UK†Additive Manufacturing and 3D Printing Research Group, Faculty of Engineering, The University of Nottingham, Nottingham, NG7 2RD, UKAbstractSelective laser melting (SLM) has shown to be an attractive manufacturing route to produce Ti-6Al-4V components. In the present study, the relation between the texture evolution and a particular set of process parameters adopted for the production of near fully dense components was investigated. The aim of this investigation was to understand which microstructural features can be tailored during the SLM production. The microstructural characterisation of the as-built components was carried out using various microscopy techniques including optical, scanning electron and transmission electron microscopy.

The microstructural characterisation of the as-built components was carried out using various microscopy techniques including optical, scanning electron and transmission electron microscopy. Phases and texture analysis were carried out using backscattered electron imaging and electron backscattered diffraction. It was found that the components consist of ’ martensitic phase precipitated from prior  columnar grains. It was found that the ’ martensitic laths develop from the parent  grains following the Burgers orientation relationship. The texture of the prior  phase was calculated and compared to the measured ’ texture. It was found that at each layer processing the ’ texture is retained and variant selection takes place. The observed microstructural details will be discussed in relation to the SLM parameters.1.

It was found that at each layer processing the ’ texture is retained and variant selection takes place. The observed microstructural details will be discussed in relation to the SLM parameters.1. IntroductionSelective laser melting (SLM) is a layer by layer additive manufacturing technique based on an infra-red fibre laser that creates solid layers from loose powder material and joins them in an additive manner. In principle, a thin layer of loose powder is initially levelled across a process platform and once the powder bed is ready selected areas of it are melted and consolidated line by line by a scanning laser. As soon as this layer is formed, new loose powder is deposited and melted on top of the layer below. Repeating this basic principle for multiple additive layers any 3D component can then be created, saving waste material and energy consumption when compared to traditional processing techniques [1].

Repeating this basic principle for multiple additive layers any 3D component can then be created, saving waste material and energy consumption when compared to traditional processing techniques [1]. As the energy delivered by the laser is fully controllable SLM has shown the possibility to fabricate near fully dense objects made of a broad range of metals generating interest within the aerospace and medical devices areas of application [2,3]. However, the complexity of the processing/microstructure and properties relationships makes SLM a manufacturing technology that has to yet be fully understood. For example, the microstructure of SLM Ti-4806Al-4V is not homogeneous due to the line- and layer-wise building strategy. The planes perpendicular to the building platform of as fabricated components always show large elongated prior  grain boundaries that are filled with acicular ’ martensitic laths as a consequence of the thermal history experienced by the layers [3-5].

The origin of this microstructure can be explained as follow: as the laser hits the powder bed the grains in the previous layers transform to the phase field. Heat is then gradually conducted away vertically through the build and the grains can solidify and grow epitaxially along the build direction in the  phase field [3,4]. As the laser moves following a pre-set scanning pattern the  grain growth freezes and ’ martensitic phase originates within the  grains due to the imposed cooling rate. This fine microstructure is responsible for the reported low ductility of the as-built components whereas the vertical prior  columnar grain boundaries have been indicated as preferential regions for crack propagation and possible source of mechanical anisotropy [3,6].

Interestingly, it has been reported that despite the fact that each layer is thermally affected by the deposition of subsequent one or more layers, the morphology and size of the ’ grains remains the same throughout the whole sample showing no dependence on the sample size [4,7]. The planes parallel to the building platform reveal the laser scan tracks and the ’ grains within prior  grain boundaries delineating the top of the prior columnar  grains. Even if the aim of SLM is to produce fully dense solid objects, typically lenticular pores oriented horizontally between two processed layers (inter-layer porosity) or rounder voids characterise all the as fabricated builds [4,7].

Numerous studies have focused on tuning several process parameters in order to develop a suitable processing window to reduce porosity: laser scan speed, hatch spacing and scan strategy have shown to play an important role in the densification mechanism of SLM parts [8-10].A systematic understanding of their contribution to the microstructure of SLM TI-6Al-4V is however still required. This study focuses on the crystallographic texture evolution of Ti-6Al- 4V produced by selective laser melting and attempts to clarify whether and which microstructural features could be possibly tailored during the building process.The aim of the present study is to understand:Texture evolution of the as fabricated components in relation to the adopted processparametersVariant selection mechanism at the basis of the /’ transformation that occurs ateach layer processing2.

Materials and Methods2.1 Powder MaterialThe starting material was a pre-alloyed plasma atomized Ti-6Al-4V grade 23. The powder was supplied by LPW, UK. In order to choose a suitable layer thickness, the powder size and distribution were initially determined using a Malvern MasterSizer 2000. The measured particle size distribution was within 15 and 70 m, but, approximately 75% of the481examined particle size ranged between 25 to 50 m. Particles appear spherical, smooth with few smaller condensed particles attached to bigger particles (Figure 1). The microstructure of the starting powders consisted of lamellar  phase with only a limited amount of retained  phase present (less than 4%). This may be due to the high cooling rate experienced by the powder during plasma atomization.

This may be due to the high cooling rate experienced by the powder during plasma atomization. The powder material microstructure is shown in Figure 2.Figure 1: Shape and morphology of the starting powder bedFigure 2: image quality of an ion-beam milled Ti-6Al-4V plasma atomised particleTable 1: Renishaw SLM250 Process ParametersLaser Power, P (W)200Hatch Spacing, h (m)100Point Distance, (m)50Exposure Time, (s)220Layer Thickenss, (m)50Scan StrategymeanderBuilding Plate Temperature, (oC)652.2 Selective Laser MeltingSimple cubic components of 1 cm3 were built using a Renishaw SLM250. The machine is equipped with a fibre modulated pulse laser, with a maximum power of 200 W. The wavelength of the laser is 1070 nm. A series of experiments were performed to establish a processing window that could minimize residual porosity and that lead to the choice of the processing parameters summarized in Table 1. A multi-directional meander scan strategy was used, i.e.

A multi-directional meander scan strategy was used, i.e. the laser scan direction was rotated by 67° at each layer in order to reduce the content of residual stresses. The cubes were built directly on top of a Ti- 6Al-4V build plate pre-heated at 65°C inside protective atmosphere back filled with Ar gas.4822.3 Optical and Electron MicroscopyThe microstructure of the as-built components was studied on specimens taken from the frontal section, i.e planes parallel to the building direction. All sample surfaces were prepared by standard metallographic preparation procedures and then etched in Kroll’s solution for optical/SEM. The microstructure of the components was examined in a Reichert-Jung MEF3 optical microscope and a Carl Zeiss (Leo) 1530 VP FEG-SEM system, which had a field emission source. Energy dispersive X-ray Spectroscopy (EDX) was used to identify the chemical composition of the powder, built components and the existence of the secondary phases.

Energy dispersive X-ray Spectroscopy (EDX) was used to identify the chemical composition of the powder, built components and the existence of the secondary phases. The distribution of crystallographic orientations of the sample crystals, i.e. the texture, was studied using electron back-scatter diffraction (EBSD) analysis available on the FEI Nova 600 Nanolab Dualbeam FIB/FEG-SEM system with EDAX EBSD camera. Data were processed using the TSL acquisition and data analysis software and MATLAB (The Mathworks, Inc., Natick, MA). In order to be able to fit the mounted samples onto the 70° pre-tilted (optimise EBSD data collection) aluminium holder, the mount was ground sideways. The scanning step size for EBSD was of 0.3m. In order to gain a further understanding of the fine microstructure, thin TEM disks were prepared. Electropolishing was carried out in a temperature range of -20 to -30oC at 20V with the A2 electrolyte supplied by Struers.

Electropolishing was carried out in a temperature range of -20 to -30oC at 20V with the A2 electrolyte supplied by Struers. The TEM analyses were performed on a JEOL 2000FX TEM operated at 200 kV.2.4 Automatic reconstruction of prior  grain orientationsSince the joining of adjacent layers occurs in the  phase field at elevated temperature, it is of interest to reconstruct the crystallographic texture of the prior  columnar grains. In this study, the reconstruction of the  grains was based on the Burgers crystallographic relationship that governs the  →  phase transformation for which the (0001) // (110) and 〈 ̅ 〉 // 〈 〉. It is reported in fact that when residual  phase is not present (as expected for as-built SLM components) the nucleation of  grains from the  laths follows the Burgers orientation relationship [11,12]. Orientations of the parent  grains were computed considering every triplet of neighboring ’ variants in the EBSD dataset.

Orientations of the parent  grains were computed considering every triplet of neighboring ’ variants in the EBSD dataset. The ’ orientation was initially expressed by a rotation matrix using the three Euler angles between the principal crystallographic axes for each ’ grain and certain axes fixed by the geometry of the specimen. Using the Burgers orientation relationship six possible parent phase ( grain) orientations were generated for every ’ belonging to the triplet. Only the solution common to the triplet (or the solution with the least misorientation) was taken as orientation of the parent  phase from which the variants had generated. Solutions exceeding a misoriention of 8° were discarded from the original dataset as a compromise to reconstruct the parent  phase with certainty and be able to generate an orientation map from a representative number of data points.4833.

Results and DiscussionTi-6Al-4V is an allotropic alloy organised in a  phase above the  transus temperature of the alloy (typically above 950oC) and in a + phase below this critical temperature. Nevertheless the amount of retained  phase at room temperature is dictated by the cooling rate experienced by the alloy and because during SLM each layer cools down very rapidly (orders of thousands degrees per second) the components can be expected to be completely ’ phase [13,14]. The optical microscopy analysis reveals in fact that the microstructure of the builds is composed of fine acicular ’ grains throughout the sample (Figure 3A) as it will be demonstrated hereafter by TEM investigations. In this study, the ’ grains do not vary with distance from the building platform or across the scanning plane. This is due to the fact that the building platform, kept at a low temperature, did not contribute to the evolution of the as deposited phase during the building process.

This is due to the fact that the building platform, kept at a low temperature, did not contribute to the evolution of the as deposited phase during the building process. It has been reported in fact for electron beam melting or laser metal deposition the builds sit at 500 to 800oC for several hours and as result of this the  grain size varies on different sample locations depending on the sample size and geometry [15,16].

This suggests the importance of the build chamber temperature in the microstructure evolution and implies that the meander scanning strategy alone cannot bias either a ’ to  decomposition or ’ grain growth.ABCFigure 3: A) Optical micrograph montage showing the microstructure of a section parallel to the building direction of as-fabricated SLM Ti-6Al-4V (prior  columnar grain boundaries indicated by the arrows); B) and C) Backscattered electron images from the same section showing the morphology and the arrangement of the ’ laths within the prior  columnar grains484The ’ martensitic phase is organized within prior  grain boundaries (Figure 3A, B and C) with different inclinations of ± 45o and 90o to the building direction as reported elsewhere [3,6].

The ’ grains do not precipitate along the  grain boundaries (Figure 3C) and this suggests that the ’ originates simultaneously from different points within the parent  grain as typical of martensitic transformations [17]. This can be explained considering the large cooling rate typical of SLM. The prior  grain boundaries are continuous with an average width of 62 ±12m. It has been reported that the width of the prior  grains is related to the hatch spacing (unitary ratio) although in this is not the case for the present study where the hatch spacing was kept at 100 m [4]. As mentioned, the presence of the prior  columnar grain boundaries is explained by the fact that Ti-6Al-4V solidifies in the  phase field and heat is mainly conducted away vertically.

As mentioned, the presence of the prior  columnar grain boundaries is explained by the fact that Ti-6Al-4V solidifies in the  phase field and heat is mainly conducted away vertically. As the scanning direction is rotated by 67o at each layer the prior  columnar grain boundaries are vertical and appear not to be skewed towards any direction in contrast to that reported for SLM Ti-6Al-4V when the scanning direction was rotated by 90o at each layer [4]. The martensitic nature of the ’ grains was confirmed studying the crystallographic structure and the chemistry of the laths under TEM. The ’ grains appear as relatively long and narrow laths characterized by high dislocation density stacking faults and twin structures (Figure 4A and 4B).Figure 4: A) Bright field TEM showing the morphology of several ’ grains, B) Typical substructure of the ’ grains; inset showing a selected area diffraction pattern relative to the ’ grain shown in Figure 4B.

The orientation of the beam was < ̅ 3 >In recent years crystallographic texture information has shown to be extremely useful and complementary to the standard microstructural studies because micro-texture plays a crucial role in the mechanical performance of several + Ti alloys [18-20]. For this reason an EBSD analysis on a frontal section of as deposited materials was carried out. The corresponding inverse pole figure orientation map (RD-TD plane) is shown in Figure 5A. This ’ texture study reveals that there isn’t a prevailing strong orientation texture (Figure 5A and C). A relatively high number of variants within each single prior  grain can be noticed (Figure 5A). Similar textures have been reported for Ti-6Al-4V processed by other AM technologies [15] although the micro-texture of as fabricated SLM doesn’t show any colony485formation of  laths sharing the same crystallographic orientation.

However, few crystal orientations tend to repeat within the same prior  grain (e.g. Figure 5A prior  grain 1 and 2), as indicated by the repetition of the same colour of the ’ grains, suggesting that texture inheritance and variant selection have taken place. In other words when the room temperature ' phase is heated in the  phase field and cooled down to room temperature again the same variants are formed [11,12].ABCDEFFigure 5: A) Orientation map of the frontal section of as-built SLM Ti-6Al-4V ’ phase and B) the reconstructed prior  grains. C) Contour and pole figures corresponding to the whole section in analysis of the ’phase and D) the reconstructed  phase. E) and F) Orientation relationship of the ’ laths belonging to two indicated prior  grains486In order to clarify this aspect, a detailed analysis of the variants within several prior  grains and their relationship with their parent  phase were analysed.

As a typical example, the orientation of the ’ laths belonging to the two labelled grains in Figure 5A and their orientation relationship with the parent phase is reported in Figure 5E and 5F. Typically 5 to 6 variants are formed inside every parent  grain and, as expected, the →’ transformation follows the Burgers orientation relationship; the (0001)’ poles are parallel to one of the (110) poles of the reconstructed parent phase and at least one of the < ̅ >’ directions is parallel to one of the <111> (Figure 5E and 5F). It was noticed that most of the analysed grains had the (0001)’ poles 60° misoriented around the < ̅ >’ axes of rotation. Only occasionally different ’ grains belonging to the same parent  grain shared a (0001)’ reflection. The ( ̅ )’ pole figures corresponding to the ’ grains with distinct (0001)’ reflections did not show any specific geometric pattern contrary to that observed for Ti-6Al- 4V fabricated by other AM technologies [21].

The reported triangular or rhombic patterns of the ( ̅ )’ pole figures are in fact a result of ’ grains with different two or three different basal poles and at least one common ( ̅ )’ pole [22,23]. This orientation relationship never occurred in the present case study. The few ’ grains sharing the (0001)’ pole had, however, the three ( ̅ )’ reflections misoriented by 10.53° around the <0001>’ axes (Figure 6). The only occasions when this happened was in correspondence to two ’ grains appearing in two different shades of the same colour (e.g. Figure 5A, grain3).

The only occasions when this happened was in correspondence to two ’ grains appearing in two different shades of the same colour (e.g. Figure 5A, grain3). This particular orientation relationship can be explained considering the 6-fold symmetry of the HCP lattice and by the fact that during the →’ transformation the < ̅ >’ direction remains parallel to the two possible <111> close packed directions that lie on the {110} plane [23].ABCDFigure 6: A) and B) (0001) and ( ̅ ) pole figures of the ’ laths of grain indicated in Figure 5B; the red arrows indicate the two variants that share a common (0001) pole and have the ( ̅ ) reflections misoriented by 10.53o; C) and D) (110) and (111) pole figures of the reconstructed prior  grain 3487In order to understand whether all the 12 ’ variants were developed with equal probability, statistics regarding the variant frequency was determined for the dataset under analysis.

The 12 ’ variants were not developed equally during cooling as shown in Figure 7. Instead, certain variants were formed more frequently than the others: the D-1C4x+,D-1C2z solutions were favoured having a frequency of 17% followed by solutions D-1C32-, D-1E and D-1C4z- (13%). Solutions D-1C31+ and D-1C4z- never occurred. The reason behind this variant selection is subject of on-going work. The fact the only a subset of the 12 different possibilities for the →’ transformation take place however is interesting because demonstrates that if the prior  grains morphology and texture is controlled during processing the final micro texture of the components can be tailored as well.Figure 7: Variant frequency distribution originated from several grains studied in Figure 5For this reason the texture of the prior  phase was calculated. Figure 5C represent the orientation map of the reconstructed  phase.

Figure 5C represent the orientation map of the reconstructed  phase. The black lines in the orientation map of Figure 5C represent high angle grain boundaries (>10°). As expected the grain boundaries of the reconstructed map match with the prior  columnar grain boundaries of Figure 5A revealing once again that columnar epitaxial growth of the  phase. The  grains texture reminds of the standard <100> solidification texture of the BCC materials, although it is slightly deviated towards the <111> texture. This result is analogous to what reported for electron beam melted Ti-6Al-4V where a strong predominance of the <100> was found [15]. Varying the scan direction at each layer seems then not to impose any particular preferred texture other than the <100> solidification texture.

Varying the scan direction at each layer seems then not to impose any particular preferred texture other than the <100> solidification texture. The correlation between the laser beam motion and the crystallographic texture of the  phase will be studied further in order to understand whether the scan strategy could be used to grow  grains with a desired direction, size and orientation.4884. ConclusionThe objective of this study was to understand the evolution of the microstructure and the texture of Ti-6Al-4V fabricated using a set of process parameters that enabled the production of near fully dense parts. This study has shown that the as-fabricated components consist entirely of ’ martensitic phase as confirmed by TEM diffraction pattern analysis and TEM- EDX. It was found that the ’ grains precipitate and replace entirely the prior parent  grains leaving behind the grain boundaries only.

It was found that the ’ grains precipitate and replace entirely the prior parent  grains leaving behind the grain boundaries only. It was demonstrated that the width of the prior  columnar grains is correlated to the hatch spacing and that rotating the scan direction at each layer SLM Ti-6Al-4V solidifies vertically and almost parallel to the building direction. Although the ’ texture of the as-fabricated components appears random due the high number of formed variants, texture inheritance and variant selection takes place. It was found that the → ’ transformation is governed by the Burgers orientation relationship: most of the ’ laths belonging to the same parent  grain have the (0001)’ reflection 60° misoriented around the < ̅ >’ axes of rotation whereas the ( ̅ )’ reflections don’t seem to follow a precise pattern.

Interestingly some of the 12 possible variants occur more frequently than others in the present dataset: this is an encouraging result for SLM of / Ti alloys because it demonstrates that if the  phase solidification texture is controlled also the ’ phase texture can be tailored. Varying the scan strategy at each layer doesn’t alter significantly the standard solidification <100> texture of the parent  phase. Further studies on the correlation of the scan strategy and the solidification texture are on-going. One of the key challenges for SLM is the possibility to tailor the microstructure during the building process: the first step was, however, to comprehend the relationship among the microstructural features of SLM Ti-6Al- 4V. Microtexture contains all the information that explains the effect of the adopted process parameters and should be considered a tool of investigation for the production of tailored SLM microstructures.4895. References1. Reeves P.

Acta Materialia 2007;55:5718-5727.12. Daymond MR, Holt RA, Cai S, Mosbrucker P, Vogel SC. Texture inheritance and variant selection through an hcp–bcc–hcp phase transformation. Acta Materialia 2010;58:4053-4066.49013. Roberts IA, Wang CJ, Esterlein R, Stanford M, Mynors DJ. A three-dimensional finite element analysis of the temperature field during laser melting of metal powders in additive layer manufacturing. International Journal of Machine Tools and Manufacture 2009;49:916- 923.14. T V, C C, JD B. As-Fabricated and Heat-Treated Microstructures of the Ti-6Al-4V Alloy Processed by Selective Laser Melting. METALLURGICAL AND MATERIALS TRANSACTIONS A 2011;42:3190-3199.15. Al-Bermani S, Blackmore M, Zhang W, Todd I. The Origin of Microstructural Diversity, Texture, and Mechanical Properties in Electron Beam Melted Ti-6Al-4V. Metallurgical and Materials Transactions A 2010;41:3422-3434.16. Facchini L, Magalini E, Robotti P, Molinari A.

Metallurgical and Materials Transactions A 2010;41:3422-3434.16. Facchini L, Magalini E, Robotti P, Molinari A. Microstructure and mechanical properties of Ti-6Al-4V produced by electron beam melting of pre-alloyed powders. Rapid Prototyping Journal 2009;15:171-178.17. Ahmed T, Rack H. Phase transformations during cooling in [alpha] [beta] titanium alloys. Materials Science and Engineering A 1998;243:206-211.18. Glavicic M, Kobryn P, Bieler T, Semiatin S. An automated method to determine the orientation of the high-temperature beta phase from measured EBSD data for the low- temperature alpha-phase in Ti-6Al-4V. Materials Science and Engineering A 2003;351:258- 264.19. Glavicic M, Kobryn P, Semiatin S. Validation of an automated EBSD method to deduce the [beta]-phase texture in Ti-6Al-4V with a colony-[alpha] microstructure. Materials Science and Engineering A 2004;385:372-376.20. Kobryn P, Semiatin S.

Materials Science and Engineering A 2004;385:372-376.20. Kobryn P, Semiatin S. Microstructure and texture evolution during solidification processing of Ti-6Al-4V. J Mater Process Technol 2003;135:330-339.21. Baufeld B, Van der Biest O, Dillien S. Texture and crystal orientation in Ti-6Al-4V builds fabricated by shaped metal deposition. Metallurgical and Materials Transactions A 2010:1-11.22. Bhattacharyya D, Viswanathan GB, Fraser HL. Crystallographic and morphological relationships between β phase and the Widmanstätten and allotriomorphic α phase at special β grain boundaries in an α/β titanium alloy. Acta Materialia 2007;55:6765-6778.23. Bhattacharyya D, Viswanathan G, Denkenberger R, Furrer D, Fraser HL. The role of crystallographic and geometrical relationships between [alpha] and [beta] phases in an [alpha]/[beta] titanium alloy. Acta materialia 2003;51:4679-4691.491.

sed Deposition Modeling of MetalsJorge Mireles1, David Espalin1, David Roberson1, Bob Zinniel2, Francisco Medina1, Ryan Wicker11. W.M. Keck Center for 3D Innovation, The University of Texas at El Paso, El Paso, TX 2. Stratasys, Inc., Eden Prairie, MNAbstractStudies have been conducted to improve previous work performed in developing a Fused Deposition Modeling for metals (FDMm) system used for applications in electronics and fabrication of 3-dimensional metallic structures. A FDM 3000 system was modified to achieve controlled deposition of eutectic Bi58Sn42 and non-eutectic Sn60Bi40 materials. Toolpath command modifications were required to achieve controlled deposition of metals. Results are presented which include a redesigned metal deposition head, computer modeling of fluid flow, and finally examples of the successful deposition of metal alloys. Additionally, FDMm- fabricated metal samples were prepared and analyzed using optical and scanning electron microscopy.

Additionally, FDMm- fabricated metal samples were prepared and analyzed using optical and scanning electron microscopy. Controlled deposition of metals using FDMm allows for parts that can be used for jigs and fixtures, electroforming mandrels, encapsulation molds, dies, electronic joining applications, as well as printing 3-dimensional electronic circuitry.1 IntroductionAdditive manufacturing (AM) allows complex parts to be built without the need for tooling, dies, or molds, using little human intervention. Fused Deposition Modeling (FDM) technology is an AM process that builds 3D shapes by taking filaments of thermoplastic polymer materials and driving them into a heated liquefier to be extruded through a small diameter nozzle onto a build platform. Technologies capable of processing metals include, for example, electron beam melting, direct metal laser sintering, laser engineered net shaping and selective laser melting.

Technologies capable of processing metals include, for example, electron beam melting, direct metal laser sintering, laser engineered net shaping and selective laser melting. Utilizing FDM technology to extrude metals poses advantages and disadvantages when compared to methods that currently build using metal alloys. An advantage of using FDM is the lack of expensive lasers equipped in sintering processes or an electron beam as is present in the electron beam melting process. Less expensive materials and systems are available that use FDM technology compared to sintering and melting technologies. A difference is also the ability to build using both thermoplastics and metals within the same build which is not possible with other direct metal systems. Disadvantages include the limitation to low-temperature, low- strength, alloys as well as the possibility for oxidation during the build process due to the lack of a controlled environment.

Disadvantages include the limitation to low-temperature, low- strength, alloys as well as the possibility for oxidation during the build process due to the lack of a controlled environment. Extending FDM to be able to process metal alloys, called FDM for metals (FDMm), has been the subject of previous work where successful deposition of tin- bismuth (Sn60Bi40 and Bi58Sn42) alloys was demonstrated (Mireles et al., 2012) . Prior research in metallic FDM with metal/polymer filaments showed successful fabrication of 3D structures (Masood and Song, 2004; Agarwala et al., 1996); however, the parts were shown to possess insufficient material properties, such as conductivity, as compared to the bulk or pure alloys (i.e. metals without a polymer binder). The research presented here utilizes a Stratasys FDM 3000 system to deposit low melting temperature (Tm) alloys (i.e. alloys of Tm below ~300oC), similar to the system described in Mireles et al., 2012.

alloys of Tm below ~300oC), similar to the system described in Mireles et al., 2012. Drawbacks in the system described in Mireles et al., 2012, include a liquefier design which presents unnecessary frictional effects due to a 90o bend and liquefier length which poses start/stop deposition issues with the836molten droplets as has been prevalent in microdispensing techniques (Bellini and Bertoldi, 2004; Bellini, 2002).The work presented here extends the concepts described in Mireles et al., 2012 to moreeffectively deposit low Tm metal alloys using a redesigned deposition head. The redesigned liquefier was conceived using current liquefier designs of machines currently produced by Stratasys (FDM Titan, FDM Maxum, 400mc, 900mc) as well as 3D printers which use the concepts of FDM and use straight liquefier configurations (MakerBot, 3DTouch). Previous work by Rice et al.

Previous work by Rice et al. 2000 has attempted to deposit low Tm metals using a modified design that implements a rheocast liquefier that uses the basic FDM process of layer-to-layer stacking with metal slurries of binary alloys instead of a wire feed process. However, the layer thickness produced using the rheocast liquefier system only achieves layer thickness on the order of 3.5mm (0.135in). Modifications of temperature, flow parameters, as well as toolpath commands have been modified to achieve controlled deposition. Flow modeling of both the redesigned and the original liquefier design are discussed with experimental results as well as possible ways to implement such designs with high temperature alloys. Metallographic images have been taken to observe the microstructure of parts built using FDMm and determine possible implications or advantages of using a FDMm system to build both electronic circuitry and 3D parts.

The applications that can benefit from low Tm metal alloys include building jigs and fixtures, electroforming mandrels, encapsulation molds, dies, electronic joining applications, as well as printing 3-Dimensional circuitry (Ojebuoboh, 1992; White and Ferriter, 1990). The following describes the design and development of the redesigned FDMm liquefier, fluid flow modeling comparing original and redesigned liquefiers, followed by deposition results, and a successful example of the fabrication of a 3D object.2 Design and conceptA FDM 3000 system (Stratasys, Inc., Eden Prairie, MN) was used in this research due toits ability to alter build parameters such as deposition head motion speed, material flow rates, chamber and liquefier temperatures, and toolpath commands. Modern FDM systems, such as the Fortus 900mc and 400mc, include a straight liquefier configuration that formed the basis for a redesign of the FDM 3000 head.

Modern FDM systems, such as the Fortus 900mc and 400mc, include a straight liquefier configuration that formed the basis for a redesign of the FDM 3000 head. Figure 1(a) shows the original configuration for a FDM 3000 deposition head and Figure 1(b) shows the redesigned liquefier implemented in this research.liquefierstepper motorsheating coilsnozzlea)Figure 1: FDM head of a) original liquefier design and b) redesigned liquefier showing difference in liquefier configuration (note that the original liquefiers is curved while the redesigned liquefier is straight as highlighted by red dashed lines)837b)The redesigned liquefier was meant to remove the 90o bend to enable the reduction of drive forces required to deliver material as well as facilitate the reduction of puddling effects due to extra material being dispensed after the stepper motors are commanded to stop – a problem inherent to the original liquefier design.

Also, the liquefier was shortened to 1) improve transient response due to the starting and stopping of the stepper motors to achieve uniform line thicknesses, 2) reduce pressure to minimize driving forces that can sometimes lead to buckling and puddling, and 3) improve heat transfer through a more concentrated heat coil configuration to achieve a smaller temperature gradient between the liquefier inlet and outlet. Through the use of a redesigned liquefier, it was hypothesized that material deposition would be better controlled in start/stop operations (i.e. material will not spill or puddle), friction effects will be reduced by incorporating a straight liquefier, and better heat transfer from the liquefier inlet to the nozzle exit.3 Experiments and methods 3.1 ModelingModeling of flow was performed with the use of ADINA software (ADINA R&D, Inc., Watertown, MA) to qualitatively demonstrate the benefits of the redesigned liquefier.

The 2D geometries were replicated for the redesigned and original liquefier versions and the same boundary conditions were applied in the modeling. The boundary conditions were laminar flow, no-slip conditions between the liquefier walls and the molten metal alloy, fluid incompressibility, gravitational acceleration of 9.8m/s2 acting on the fluid, steady-state conditions, (Roxas and Ju, 2008, Bellini, 2002), inlet velocity of 3.73mm/s (0.147in/s) (as measured), and the material properties for molten Sn-Bi alloy of 15,000cP (as measured for molten alloy using a Brookfield DV-E Viscometer).3.2 Experiments and methodsExperimentation performed on the fluid flow was qualitative, as instrumentation tomeasure pressure was unavailable, and results included observations of material flow through each nozzle.

To build complex 2D shapes as well as facilitate the stacking of layers, modifiable toolpath commands were generated using Insight software (Stratasys Inc., Eden Prairie, MN) from a corresponding CAD file. A modified tip was also used which has an increased nozzle diameter of 0.635mm (0.025in) that allows for improved deposition of molten metal and decreases pressure within the liquefier allowing the filament to be fed with less applied force by the servo motors. During the build process, the layer thickness (deposition head movement in Z- direction) was modified to account for increased deposition from the modified tip to achieve a layer thickness of 0.762mm (0.03in). Other build parameters changed include the slowdown of the deposition head velocity, the flow, and the liquefier temperature. The deposition head velocity was decreased to 750microsteps per second from 1000microsteps per second.

The deposition head velocity was decreased to 750microsteps per second from 1000microsteps per second. A full explanation of the flow parameter modification is found elsewhere (Mireles et al., 2012).Additionally, experiments involving the variation of thermal attributes were performedusing type K thermocouples. The liquefier temperature was set to 220oC through the temperature controllers in the FDM 3000. The envelope temperature was set to ambient temperature and the neighboring liquefier was set to a lower temperature (210oC) and the tip was removed to avoid excess heat transfer between the model and support liquefiers. A thermocouple was placed at the nozzle tip and the temperature was read once the system reached838steady temperature conditions. The liquefier temperature was set to 210oC for the eutectic alloy Bi58Sn42 and 220oC for the non-eutectic alloy Sn60Bi40.

The liquefier temperature was set to 210oC for the eutectic alloy Bi58Sn42 and 220oC for the non-eutectic alloy Sn60Bi40. These temperatures account for a temperature difference from the heat supplied to the liquefier by the heating coils to the heat conducted to the nozzle end.Scanning electron microscopy (SEM) was performed with a Hitachi S-4800 Ultra-highResolution Field Emission Scanning Electron Microscope (Hitachi High-Technologies Corporation, Tokyo, Japan) equipped with an EDAX energy dispersive X-ray analyzer (EDS) and utilizing a 20keV accelerating voltage. Optical metallography was performed using a Leica MEF4M optical digital imaging system on a mounted sample of stacked layers built using the FDMm system. Metallography was performed on the mounted samples using 400grit, 1000grit, and 1200grit paper, followed by a finishing step with a polishing cloth using 0.3µ alumina slurry.

Metallography was performed on the mounted samples using 400grit, 1000grit, and 1200grit paper, followed by a finishing step with a polishing cloth using 0.3µ alumina slurry. An etchant composed of 10% HCl and 90%H2O was applied to reveal the microstructure.4 Results 4.1 Modeling resultsFigure 2 shows the flow modeling and pressure results for the original liquefier design(Figure 2(a)) and the redesigned liquefier (Figure 2(b)). Pressure was of importance since it directly affects how the material is manipulated during deposition (i.e. a greater pressure drop increases the amount of force needed to push the material filament to achieve extrusion) (Bellini, 2002; Roxas and Ju, 2008). The pressure drop between both designs differed and was ~14% greater in the original liquefier design.

The pressure drop between both designs differed and was ~14% greater in the original liquefier design. The same modeling methodology was also performed in different research where a curve liquefier was modeled (Ramanath et al., 2008); separate work modeled a straight liquefier (Bellini, 2002) and both showed noteworthy differences between designs. During metal deposition experiments, the start and stop behavior of flow was observed and showed continued flow after the filament was no longer fed using the original liquefier. The continued flow was not observed when using the redesigned liquefier and may be due to the difference in liquefier length and the liquidus properties of the metal alloy inside the liquefier.

The continued flow was not observed when using the redesigned liquefier and may be due to the difference in liquefier length and the liquidus properties of the metal alloy inside the liquefier. The amount of force needed to push the filament to facilitate continuous deposition was also reduced 150% from previous research results (Mireles et al., 2012) which was likely attributed toa)Figure 2: Pressure gradient for a) original liquefier design and b) redesigned liquefier839b)6.35mmb)a)Figure 3: Demonstration of deposition results using FDMm where a) shows an example of 2D shapes and b) shows a 3D structure of 6 stacked layers.lowering the pressure with the redesigned liquefier.

Additionally, experimentation performed utilizing type K thermocouples showed a difference between the temperature set by the operator (220oC) to the temperature at the nozzle of 65oC for the original liquefier design compared to 48oC for the redesigned liquefier demonstrating improved precision.Using ideal conditions, the simulation results for pressure still showed a notabledifference between both designs. The non-ideal conditions including frictional effects, slip between the liquefier walls and the filament, as well as uneven heat distribution, can produce a greater difference between the two liquefier designs and demonstrates the need for a straight liquefier design as this design minimizes the non-ideal conditions.

Further emphasizing the need for a straight liquefier design is the fact that similar designs are already used in newer FDM models.4.2 Deposition resultsTo demonstrate the improved capability of the redesigned liquefier as well as the effect of code modifications and liquefier temperature changes, 2D and 3D structures were fabricated and are shown in Figure 3(a) and Figure 3(b), respectively. As shown, layer stacking was achieved and parts were created using metal alloys of low Tm. Road width of the layers using eutectic material was 1.24mm (0.049in) and layer thickness was 0.74mm (0.029in). For non-eutectic material, road width was 1.12mm (0.44in) and layer thickness was 0.71mm (0.028in).

As shown in Figure 3, the lines for both 2D and 3D parts are continuous throughout. It was previously hypothesized that non-eutectic materials are more suitable for FDM due to the presence of a “mushy” zone which allows for a higher viscosity and therefore deposition was more controlled and continuous (Mireles et al., 2012, White and Ferriter, 1990). The results in Table 1 showed little difference between both eutectic and non-eutectic materials, however, non-eutectic lines still showed less variance in thickness throughout the length of a deposited line.

Although non-eutectic alloys still perform better than eutectic compositions, the redesigned liquefier improves the deposition of eutectic alloys.conductive materialsubstrateb)c)d)a)0.6mme)Figure 4 3D electronics using FDM where a) shows the stacking of ABS vias connecting each other with the Sn-Bi conductive material, b) showing the 3D model of the conductive material within the ABS substrate, c) is the top view of a built sample, d) is the bottom view of a built sample, and e) is the magnified view of a circuit cross-section at 72X showing the ABS layers with Sn-Bi within alternate layers.841Figure 5: (a) Optical micrograph showing the eutectic microstructure and (b) corresponding phase diagram indicating the composition of our alloy (Kattner, 2002).A potential application for FDMm is the building of 3D circuitry that allows for electricalcomponents to be connected on different faces of a solid.

Figure 4 demonstrates the ability for the support liquefier in the deposition head shown in Figure 1 to deposit a substrate and for the model liquefier to deposit the conductive material that connects to each layer through vias in the substrate material. Figures 4(a) and 4(b) show the CAD demonstration of the model while Figures 4(c) and 4(d) represent the actual deposition results. Figure 4(e) shows a magnified view taken with a stereomicroscope magnified at 72X of two subsequent layers showing Sn-Bi material filling the ABS via. Figure 4 only showed contacts between layers without electronic components; however, cavities which allow for component placement can be produced as demonstrated by Lopes et al. 2012 where stereolithography was used in the creation of 3D structural electronics.4.3 Microstructural analysisThe optical metallograph (Figure 5) of the cross section of the FDM-built Bi58Sn42solder component displays a two phase eutectic microstructure.

This was consistent with what would be expected from the Sn-Bi phase diagram seen in Figure 5(b). Also observed in the crossFigure 6: (a) The layering of fine and course microstructure observed on the cross-section of the FDM-Built component and (b) higher magnification image of the same area.842Figure 7: SEM micrograph and corresponding EDS spectra showing the differences between the two phases.section (Figure 6) is a layering effect of fine and coarse microstructure which was most-likely due to the build process.The SEM micrograph (Figure 7) of cross section of the FDM-built eutectic Bi58Sn42 solder component again showed a eutectic microstructure. Spot EDS analysis of the two phases reveals a Bi-rich phase with a Sn-Bi phase interdispersed throughout. Also observed within the microstructure was the same layering of fine and coarse lamella (Figure 8). The topography observed in Figure 8 was due to etch preparation as part of the metallographic process.

The topography observed in Figure 8 was due to etch preparation as part of the metallographic process. The coarsening of the lamella may be due to the reheating of the previous build layer by the subsequent buildFigure 8: SEM micrograph showing the layering between fine and coarse microstructure843layer. In the context of a structural part, the coarsening of the eutectic microstructrure has been observed to improve the mechanical properties of Sn-Bi alloys used in various applications of low Tm alloys (Felton et al. 1993).

1993). In the context of the electric circuit presented here the coarsening may also be beneficial as, in general, a metallic conductive path with large grains will have better conductivity than the same material with small grains as grain boundaries act as detractors to the flow of electrons (Roberson, 2012).ConclusionsThrough the use of a FDM 3000 system, controlled deposition of metal alloys has beenachieved and 3D structures have been built in a layered fashion using a redesigned liquefier. Idealistic modeling was done to evaluate the performance difference between the original liquefier and the redesigned liquefier. Assuming ideal conditions there were notable differences that exemplify the need for an improved liquefier design. Experimental measurements and deposition results further demonstrated the superior performance of the redesigned liquefier over the original liquefier.

Experimental measurements and deposition results further demonstrated the superior performance of the redesigned liquefier over the original liquefier. Various applications can benefit from metallic FDM including building jigs and fixtures, electroforming mandrels, encapsulation molds, dies, electronic joining applications, as well as printing 3-Dimensional circuitry. Microstructural analysis demonstrated good interlayer bonding with variations in the coarseness of microstructure along the layer interface that affects mechanical and conductive properties. The results of this work give potential to utilize the redesigned FDMm configuration with different deposition head materials to build metallic components using higher strength alloys.Future WorkFurther work can be extended to provide results using higher Tm metals which have potential applications in the fabrication of structural parts using the same methodology explored in this research.

Accuracy has improved with the implementation of a redesigned deposition head; however, improvements are required to achieve accuracies equivalent to those achieved by polymers such as ABS (+-0.005). The fabrication of more complex 3D structures needs to be demonstrated and the microstructure and interfacial phenomena between layers for such structures needs to be analyzed. For 3D circuits, a fully-functional circuit needs to be demonstrated with cavities for electronic component placement. Control of microstructure may be accomplished by close control of temperature parameters to obtain favorable microstructures (coarsening) for maximum mechanical properties as well as uniform conductivity.AcknowledgementsThe research described in this paper was performed within the W.M. Keck Center for 3DInnovation at the University of Texas at El Paso (UTEP).

Keck Center for 3DInnovation at the University of Texas at El Paso (UTEP). Additional support was provided by the UTEP Louise Stokes Alliance for Minority Participation program funded through grant number HRD-1139929 from the National Science Foundation. The findings and opinions presented in this paper are those of the authors and do not necessarily reflect those of the sponsors of this research.844ReferencesAgarwala, M. K., Weeren, R. V., Bandyopadhyay, A., Whalen, P. J., Safari, A., and Danforth, S. C.,(1996), Fused Deposition of ceramics and metals: an overview. Proceedings of the Solid Freeform Fabrication Symposium, Austin, Texas.Bellini A., and Bertoldi M., (2004). Liquefier dynamics in fused deposition modeling,” Journal ofManufacturing Science and Engineering, 126, pp. 237-246.Bellini A., (2002). Fused deposition of ceramics: A comprehensive experimental, analytical andcomputational study of material behavior, fabrication process and equipment design. Ph.D.

sed Deposition Modeling of Polymethylmethacrylate for Use in Patient- Specific Reconstructive SurgeryDavid Espalina, Karina Arcautea, David Rodrigueza, Frank Medinaa, Matthew Posnerb, Ryan Wickeraa The University of Texas at El Paso, W.M. Keck Center for 3D Innovation, College of Engineering, El Paso, TX 79968-0521 bWilliam Beaumont Army Medical CenterAbstractPolymethylmethacrylate (PMMA) is a biocompatible alloplastic material used in craniofacial reconstruction and as bone cement and antibiotic-impregnated spacers in orthopaedics. The polymerization of PMMA in-situ causes tissue necrosis and other complications due to the long surgical times associated with mixing and shaping the PMMA. PMMA is a thermoplastic acrylic resin suitable for extrusion in FDM thus 3D anatomical models can be fabricated prior to surgery directly from medical imaging data.

PMMA is a thermoplastic acrylic resin suitable for extrusion in FDM thus 3D anatomical models can be fabricated prior to surgery directly from medical imaging data. The building parameters required for successful FDM fabrication with medical-grade PMMA filament (1/16”Ø) were developed using an FDM 3000. It was found that a liquefier and envelope temperature of 235ºC and 55ºC, respectively, as well as increasing the model feed rate by 60%, were necessary to properly and consistently extrude the PMMA filament. Scaffolds with different porosities and fabrication conditions (tip wipe frequency and layer orientation) were produced, and their compressive mechanical properties were examined. Results show that both the tip wipe frequency (1 wipe every layer or 1 wipe every 10 layers) and layer orientation (transverse or axial with respect to the applied compressive load) used to fabricate the scaffolds, as well as the porosity of the scaffold had an effect on the mechanical properties.

The samples fabricated with the high tip frequency had a larger compressive strength and modulus (Compressive strength: 16 ± 0.97 vs. 13 ± 0.71 MPa, Modulus: 370 ± 14 vs. 313 ± 29 MPa, for samples fabricated in the transverse orientation with 1 tip wipe per layer or 1 tip wipe per 10 layers, respectively). Also, the samples fabricated in the transverse orientation had a larger compressive strength and modulus than the ones fabricated in the axial orientation (Compressive strength: 16±0.97 vs. 13±0.83 MPa, Modulus: 370±14 vs. 281±22 MPa, for samples fabricated with 1 tip wipe per layer, in the transverse and axial orientation, respectively). Overall, the compressive strain for the samples fabricated with the four different conditions ranged from 8 – 12%.

Overall, the compressive strain for the samples fabricated with the four different conditions ranged from 8 – 12%. In regards to the porosity of the samples, in general, the stiffness, yield strength and yield strain decreased when the porosity increased (Compressive strength: 12±0.71 – 7±0.95 MPa, Modulus: 248±10 – 165±16 MPa, Strain: 7±1.5 – 5±1% for samples with a porosity ranging from 55 – 70%). The successful FDM fabrication of patient-specific, 3D PMMA implants with varying densities, including the model of a structure to repair a cranial defect and the model of a femur, was demonstrated. This work shows that customized structures with varying porosities to achieve tailored properties can be designed and directly fabricated using FDM and PMMA.1. IntroductionFabricating reconstructive structures has been accomplished with many techniques and technologies including traditional methods such as solvent casting, fiber-bonding, membrane lamination, melt molding, and gas forming.

While these methods have been adopted in the medical industry, they do not provide structures with consistent interconnected pore networks and uniform microstructures [Zein et al., 2002]. Reconstructive structures are developed and fabricated to play very specific roles including temporary bone spacers, antibiotic delivery devices, permanent bone substitutes, biodegradable bone structures, etc. Reconstructive structures are necessary in correcting bone injuries due to trauma, bone tissue loss after tumor surgery, and congenital bone malformations, or in providing a temporary spacer in the case of bone implant failure due to infection [Eufinger et al.

1995; Materialise, 2009; Lee et al., 2009; Frutos Cabanillas et al., 2000, Pelletier et al., 2009].Polymethylmethacrylate (PMMA) is an FDA approved synthetic biomaterial widely used to fabricate reconstructive structures, including dental implants, reconstruction of craniofacial defects, or as bone cement to remodel lost bone and affix implants. Commercially available PMMA bone cements include powder constituents of prepolymerized PMMA, an initiator that catalyzes the polymerization process and a radiopacifier. Liquid constituents contain large amounts of monomer (MMA), an accelerator that breaks down the catalyst, and a stabilizer that impedes premature polymerization [Lewis, 1997].PMMA is commonly used in orthopedic surgery in the form of the two-component system described previously. Mixing of both components is conducted and inserted into the surgery site where polymerization occurs. There are several main drawbacks with the technique.

Mixing of both components is conducted and inserted into the surgery site where polymerization occurs. There are several main drawbacks with the technique. Polymerization of PMMA in-situ causes tissue necrosis, weakened local blood circulation and tendency to membrane formation all of which are associated with the high exothermic temperature, ranging from 67°C to 124°C, of the bone cement. Also, bone necrosis is caused by unreacted MMA monomer before complete polymerization. Another complication is the size reduction of the bone cement during the polymerization process. There is also a significant stiffness difference between PMMA and the contiguous bone and risk of infection due to the large operating times [Lewis, 1997; Eufinger et al.

There is also a significant stiffness difference between PMMA and the contiguous bone and risk of infection due to the large operating times [Lewis, 1997; Eufinger et al. 1995].In order to avoid the drawbacks of polymerizing PMMA in-situ, other processing methods for PMMA have been attempted, including casting and compression molding in a cavity mold created from the master implant made in an additive manufacturing (AM) or rapid prototyping (RP) technology from medical imaging data prior to implantation surgery [Gopakumar, 2004; Lee et al., 2009, Materialise, 2009]. The molded PMMA implant can be sterilized by gas or radiation sterilization or by autoclaving. Some other researchers have used RP models of the bone defect to model the implant prior or in the surgery site [Lee et al., 2009, Materialise, 2009].PMMA is a suitable material to use in fused deposition modeling (FDM).

Fused deposition modeling is an additive manufacturing (AM) technology that allows fabrication of complex three-dimensional geometries layer-by-layer. A filament of thermoplastic material is driven at a controlled rate into a liquefier that is heated to a prescribed temperature associatedwith the material being used. The material is extruded in a semi-molten state through a small diameter nozzle while the head scans a building stage along a programmed toolpath. Upon completion of the layer, the stage lowers a predefined Z distance and successive layers are created in the same fashion. PMMA has a melting temperature range (200°C to 230°C) [Osswald et al., 2006] that lies within FDM operating temperatures (up to 290°C). Another advantage is that PMMA has a low water absorption capability (0.35%), which inhibits moisture collection that can cause problems in the FDM extrusion process.

Another advantage is that PMMA has a low water absorption capability (0.35%), which inhibits moisture collection that can cause problems in the FDM extrusion process. The characteristics of parts fabricated in FDM are influenced by material properties and processing parameters that include extrusion/envelope temperature, material feed rate, slice height (SH), road width (RW), raster angle (RA), deposition layer orientation, air gap (AG) and contour styles.This study initiates an effort at processing PMMA via FDM for medical applications including the creation of non-load bearing reconstructive bone structures and temporary bone spacers in an attempt to improve the current state of the use of PMMA in these applications. The effect of different fabrication conditions on mechanical properties and porosity of PMMA parts fabricated in FDM, as well as the relation between mechanical properties and porosity of PMMA parts were examined.2. Materials and methods2.1.

Materials and methods2.1. FDM system optimizationAdapting the FDM 3000 system (Stratasys Inc., Eden Prairie, MN) to properly and consistently extrude PMMA filament (Biogeneral Inc., San Diego, CA) required identifying parameters that dictate extrusion regardless of material and determining the suitable parameter values for PMMA. Extrusion with FDM primarily depends on liquefier temperature, envelope temperature and material feed rate. Simple square prisms were fabricated with T16 tips, an envelope temperature of 75°C and various liquefier temperatures while closely monitoring the building process. Relatively low temperatures, 180°C - 210°C, failed to properly extrude PMMA as the filament would audibly break before entering the liquefiers. A temperature of 225°C allowed PMMA extrusion but created extremely stiff rasters that would not properly adhere to the support material (ABS P400 R) resulting in models detaching before proper completion.

Higher temperatures of about 270°C created rasters with visible wavy surfaces and left material residue that affected geometry. Additionally, excessive material flowed from the extruding tips while the FDM system was not building. A liquefier temperature of 235°C was found to have the best raster surfaces and the minimum evidence of material residue for this envelope temperature. Various envelope temperatures were tested and it was found that 55°C was optimal in minimizing material residue and enhancing raster surfaces. Material feed rates were determined by executing the default calibration method supplied by the Insight V3.5 slicing software (Stratasys Inc., Eden Prairie, MN) until proper raster widths were achieved.2.2. Experimental designTwo experimental designs were carried out to determine the effect of different factors (fabrication conditions and scaffold characteristics) on the mechanical properties of the fabricated structures as described in Table 1.

Experimental Design I analyzed the effect of tipTable 1. Design of experimentsConditionsSample Criteriatip wipe frequency1 tip wipe/layer 1 tip wipe/10 layersExperiment Ianisotropytransverse axialExperiment IIporosity55% 63% 66% 70%wipe frequency and anisotropy on the mechanical properties of the fabricated structures, while Experimental Design II evaluated the mechanical properties of samples with different porosity percentages. Each experimental condition consisted of 6 square prism specimens with design dimensions 7 x 7 x 13.5 mm. The scaffold structures were designed by maintaining a 0.406mm RW, 0.254mm SH and 0°/90° RA. AG was varied in order to achieve different porosity ranges. Slicing and toolpath parameters are illustrated in Figure 1.2.2.1. Tip wipe frequencyFDM systems constantly clean the extruding tips by wiping them against a brass brush or a rubber.

Slicing and toolpath parameters are illustrated in Figure 1.2.2.1. Tip wipe frequencyFDM systems constantly clean the extruding tips by wiping them against a brass brush or a rubber. When a part is manufactured, the parameters for tip wipe configuration can be set at different values to produce cleaner parts or to save time and material. In the default configuration the system does a tip wipe after depositing ten layers (1wipe per 10 layers). A tip wipe every deposited layer (1 wipe per layer) is usually configured in order to produce cleaner parts with less contamination or blend of materials between layers. For Experimental Design I, two different tip wipe frequencies were tested: 1 wipe every ten layers, and 1 wipe every layer.2.2.2.

For Experimental Design I, two different tip wipe frequencies were tested: 1 wipe every ten layers, and 1 wipe every layer.2.2.2. Layer orientationTwo different layer orientations were tested for Experimental Design I, with one having layer orientations in the transverse position relative to applied force and the other orientated in the axial position as illustrated in Figure 2. These two layer orientations were selected to determine anisotropic influences on mechanical properties of the samples.2.2.3. Percent porosityA range of porosities within the 50-70% range was achieved by varying the AG in the scaffold design. The effect of the scaffold percent porosity on the mechanical properties of the structures was evaluated in Experimental Design II.abFigure 1. Slicing and toolpath parameters (a) trimetric view illustrating road width (RW), air gap (AG) and slice height (SH), (b) top view with 0°/90° RA.2.3.

Slicing and toolpath parameters (a) trimetric view illustrating road width (RW), air gap (AG) and slice height (SH), (b) top view with 0°/90° RA.2.3. Scaffold characterizationThe fabricated scaffolds were characterized in terms of their mechanical properties and porosity. The samples were tested in compression to evaluate their yield strength, yield strain and modulus of elasticity. The samples were weighted and their dimensions were evaluated to estimate the porosity. Furthermore, the specimens for Experimental Design II were imaged in order to measure the RW, AG and SH of the fabricated scaffold, and compare the measured values versus the design values.2.3.1. PorosityPorosity estimates were achieved by correlating scaffold geometrical dimensions and material density as per ASTM F 2450-04. The scaffold volume, VT , was determined from caliper measurements of length, width, and height.

The scaffold volume, VT , was determined from caliper measurements of length, width, and height. Utilizing the material density of 1.17g/cm3 [Osswald et al., 2006] the following equation was employed:Vp = VT −ms ρswhere VP is the total volume occupied by the network of pores, ms is the total specimen mass determined using a Sartorius CP124S micro-balance (Sartorius Stedium SUS Inc., Concord, CA) and ρs is the material density. Percentage of porosity was calculated according to:Porosity =VP VT×100ForceForceabFigure 2. Layer orientations relative to compressive load (a) transverse (b) axial.2.3.2. Compressive testingCompressive tests were conducted following guidelines provided by ASTM F 451-08 on an Instron 5866 system (Instron®, Norwood, MA). A 10 kN load cell was used and a deformation cross head speed of 1 mm/min was employed. Compressive stress calculations were carried out by dividing the applied force by the original specimen cross-sectional area.

Compressive stress calculations were carried out by dividing the applied force by the original specimen cross-sectional area. Strain was defined as the ratio of axial deformation to original specimen height. Compressive stress- strain graphs were created from which several mechanical properties were acquired. Compressive yield strength was defined as the stress after which the initial linear region deviated from linearity; yield strain is the strain associated with the compressive yield strength. The stiffness is defined as the modulus of elasticity, E, and was calculated from data representing the slope of the initial linear region.2.3.3. Imagingimaged using a stereomicroscope (Leica MZ16, Leica Microsystems Inc., Bannockburn, IL) equipped with a CCD camera (Retiga 2000R Fast 1394, QImaging Corp., Canada). Measurements for the RW, AG and SH were taken from digital images.

Measurements for the RW, AG and SH were taken from digital images. Several images were captured in order to take at least five measurements per dimension per sample.Specimens for Experimental Design II were observed and2.4. Statistical analysisFor all experiments, student’s t-tests were performed to compare mean values of independent samples with a significance level of 0.05. In addition, an ANOVA test was performed for the specimens used in Experimental Design II to compare RW mean values of six independent samples with a significance level of 0.05. Data in graphs represents averages ± one standard deviation.3. Results and discussion3.1. Experimental design IThe results on mechanical properties and porosity for samples built with different tip wipe frequencies and layer orientations are summarized graphically in Figure 3. Also, Table 2 contains physical data (mass, volume, and estimated porosity) for the samples.

Also, Table 2 contains physical data (mass, volume, and estimated porosity) for the samples. It is important to note that contraction and inner stresses develop in the recently deposited layer as it cools from the extruded temperature to the envelope temperature [Wang et al., 2007]. If the envelope temperature lies within the extruded and material glass-transition temperature range, stress will not accumulate amongst layers, therefore inhibiting the structure from withstanding higher stresses. The research presented here allowed the envelope temperature to be lower than the PMMA glass-transition temperature of 130°C to promote higher stress resistance.

Experiment I compressive test results (a) percent porosity, (b) stiffness, (c) yield strength, (d) yield strain.3.1.1. PorosityPorosities and physical properties are reported in Table 2. In general, VT, VP, and porosity were higher when fabricating with one tip wipe per 10 layers in both the transverse and axial orientations. In the transverse orientation, porosity increased from 48% to 56% when the frequency of tip wipes was reduced from once per layer to once per 10 layers. In the axial orientation, porosity also increased from 55% to 60 % when tip wipes were used less frequently. Student’s t-test results indicated that the effect of tip wipes was statistically significant as p ≤ 0.05 for both transverse and axial orientations.

Student’s t-test results indicated that the effect of tip wipes was statistically significant as p ≤ 0.05 for both transverse and axial orientations. Samples fabricated with one tip wipe per 10layers had less elapsed time between layer deposition that resulted in minimized contraction and promoted larger VT, VP and percent porosity.Primary differences in mass originated from more model volume required for the transverse position that inherently decreasing porosity due to layer orientation (student’s t-test, p ≤ 0.05). Additionally, each layer required a tool start and stop operation which resulted in residue material that unintentionally increased the scaffold mass. Since the transverse layer position required 54 layers compared to 28 axial layers, more residue material was collected that simultaneously increased the mass and decreased the porosity as illustrated in Figure 3a.

Therefore, the axial layer position achieved 55% porosity and the transverse position resulted in 48% porosity.Table 2. Experiment I porosities and physical properties.TransverseAxial1 tip wipe per layer1 tip wipe per 10 layers1 tip wipe per layer1 tip wipe per 10 layersAverage mass (g) Average VT (mm3) Average VP (mm3)0.38130.6300.3040.3780.7320.4090.32710.6230.3440.32820.7090.429Average Porosity (%)485655603.1.2. Compressive yield strengthResults reported in Figure 3c found that yield strength decreased from 16 MPa to 13 MPa when decreasing the frequency of tip wipes in the transverse orientation. Similarly, yield strength decreased from 13 MPa to 11 MPa in the axial orientation when tip wipe frequency was decreased. Allowing the recently deposited layers to cool longer when using one tip wipe per layer increased the inner stresses and promoted higher yield strength.

Allowing the recently deposited layers to cool longer when using one tip wipe per layer increased the inner stresses and promoted higher yield strength. In general, yield strength decreased when tip wipe frequency was decreased from once per layer to once every 10 layers (student’s t-test, ). Layers positioned in the axial direction provided columns of PMMA in every other layer to support the applied force during compressive testing. In this particular position, 14 layers supported the force whereas the transverse position provided 54 layers to resist the compressive load. Figure 3c illustrates that yield strength was higher (student’s t-test, p ≤ 0.05) in the transverse orientation than in the axial orientation.05.0≥p3.1.3. Compressive yield strainYield strain shown in Figure 3d experienced a decrease from 9% to 8% when the tip frequency was reduced from 1 tip wipe per layer to 1 tip wipe per 10 layers in the transverse orientation.

Student’s t-test supported the significance, p ≤ 0.05. The opposite was graphically noted for the axial orientation as yield strain seemingly increased from 9% to 10% when tip wipe frequency was decreased. However, student’s t-test results indicated that the effect of tip wipe configuration for the axial orientation was not significant,05.0≥p.Transverse and axial orientations experienced a yield strain of ~9% and there is evidence that the means are not different (t-test, p ≥ 0.05). This indicates that compressive tests yielded identical strains for both layer orientations but a higher yield stress was supported by the transverse layer orientation.3.1.4. StiffnessStiffness was observed to be higher when tip wipes were executed after each layer. In the transverse orientation, stiffness decreased from 370 MPa to 313 MPa when tip wipe frequency was decreased from once per layer to once every 10 layers.

In the transverse orientation, stiffness decreased from 370 MPa to 313 MPa when tip wipe frequency was decreased from once per layer to once every 10 layers. In similar fashion, stiffness decreased from 281 MPa to 202 MPa when tip wipe frequency was decreased in the axial orientation. This indicates that the cooling and contraction effect caused by frequent tip wipes allowed and created higher stiffness desirable for bone reconstructive structures (student’s t-test, 05.0≥p ). Caution must be taken as larger structures may experience significant contraction and complications may arise, therefore yielding non-conforming scaffold dimensions.Recalling the yield strength and strain results mentioned above, it is expected to find higher stiffness for specimens fabricated in the transverse position.

Results supported this hypothesis as stiffness for specimens in the transverse position was calculated to be 370 MPa (axial orientation yielded 281 MPa) with the effect of layer orientation being statistically significant (student’s t-test, p ≥ 0.05).3.2. Experimental design IIPorosity and compressive testing results are illustrated in Figure 4. Table 3 shows the measured dimensions for RW. SH and AG for the samples. After analyzing the 0.610 mm and 0.800 mm AG samples, it was concluded that there was no significant difference in porosity.3.2.1. Dimensional characteristicsStereomicroscope images were analyzed to determine RW, AG and SH. An ANOVA analysis was done on all samples and it was determined that road width means did not differ ( p ≥ 0.05) when values of AG where increased. Actual road width for all samples in Experiment II was determined to be 0.352 ±0.039. The expected 0.406 mm RW was not achieved but the difference was minimal.

Actual road width for all samples in Experiment II was determined to be 0.352 ±0.039. The expected 0.406 mm RW was not achieved but the difference was minimal. This difference in RW also caused a slight difference in AG with the largest difference occurring when the actual AG was measured as 2.645 mm and the expected AG was 2.800 mm.Table 3.

Experiment II: porosity and compressive testing results (p<0.05 when compared to “Q” 0.610 mm AG sample, “R” 0.800 mm AG sample, “S” 1.200 mm AG sample, “T” 1.700 mm AG sample, “V” 55% porosity sample, “W” 63% porosity sample, * indicates p<<0.05) (a) percent porosity, (b) stiffness, (c) yield strength, (d) yield strain.Slice heights remained relatively constant when varying AG except in the case of the 70% porosity sample. Careful observations during fabrication revealed that the expected 2.800 mm AG provided insufficient layer-to-layer foundations for deposition, therefore causing the suspended rasters to have a larger slice height. In addition, portion of the layers often collapsed to preceding layers affecting the mechanical properties.3.2.2. PorosityAs illustrated in Figure 4a, porosity increased as AG was increased resulting in samples of the following percent porosities: 56%, 55%, 63%, 66%, and 70%.

PorosityAs illustrated in Figure 4a, porosity increased as AG was increased resulting in samples of the following percent porosities: 56%, 55%, 63%, 66%, and 70%. This increase is due to VT remaining relatively constant as ms decreased and VP increased when AG increased. Another approach is to associate porosity with the number of rasters required to construct each layer. The least porous specimen is associated with the most number of required rasters, 7 per layer, while the most porous specimen contained 3 rasters per layer. All other specimens follow the same trend as shown in Figure 5.3.2.3. Compressive yield strengthMajor load support was provided at raster overlapping locations that are parallel throughout the entire specimen height as illustrated by shaded areas in Figure 5. Additionalsupport was provided by the intersecting rasters. A general trend of decreasing yield strength was observed when porosity was increased as illustrated in Figure 4c.

Additionalsupport was provided by the intersecting rasters. A general trend of decreasing yield strength was observed when porosity was increased as illustrated in Figure 4c. This trend held true at all porosities except at the highest porosity achieved, 70%. Layer segments that collapsed onto preceding layers provided added support that allowed higher stress resistance. The faulty fabrication of this sample caused the yield strength, yield strain and stiffness to deviate from the general trend. Excluding the 70% porosity sample, yield strength decreased from 12 MPa to 7 MPa when porosity increased from 55% to 66%.7.1X7.1Xad7.1X7.1Xbe7.1XcFigure 5.

Excluding the 70% porosity sample, yield strength decreased from 12 MPa to 7 MPa when porosity increased from 55% to 66%.7.1X7.1Xad7.1X7.1Xbe7.1XcFigure 5. Experiment II: samples of varying porosities with associated overlapping rasters (a) 56% porosity, 0.610 mm AG, 7 rasters per layer, (b) 55% porosity, 0.800 mm AG, 6 rasters per layer, (c) 63% porosity, 1.200 mm AG, 5 rasters per layer, (d) 66% porosity, 1.700 mm AG, 4 rasters per layer, (e) 70% porosity, 2.800 mm AG, 3 rasters per layer. Each sample is accompanied with a representation of overlapping rasters.3.2.4. Compressive yield strainYield strain shown in Figure 4d also followed a decreasing trend when porosity increased with the exception being the 66% and 70% porosity samples. Student’s t-test indicated that the 63% and 66% porosity sample means are not different ( p ≥ 0.05). The yield strain anomaly in the 70% porosity sample is also a result of the collapsed layers.

Student’s t-test indicated that the 63% and 66% porosity sample means are not different ( p ≥ 0.05). The yield strain anomaly in the 70% porosity sample is also a result of the collapsed layers. Yield strain decreased from 7% to 5% (excluding the 70% porosity sample) when porosity increased from 55% to 66%.3.2.5. StiffnessExcluding the 70% porosity sample, stiffness decreased from 248 MPa to 165 MPa when porosity increased from 55% to 66%. This trend is reflective of the trends established by the yield strength and strain. Although the application of this scaffold fabricating process is for non- load bearing structures, higher stiffness is still desirable which was exhibited by the least porous sample.3.4. Application fabrication of patient-specific 3D models3D CAD models can be generated from medical imaging data sets acquired through CT scans or MRI by utilizing image processing software such as Mimics® (Materialise NV, Irvine, CA).

CAD models can then be tailored to create scaffolds with specific porosities with FDM technology. Figures 6 and 7 show patient-specific, anatomical 3D PMMA models fabricated in a FDM 3000 system, including the model of a cranial defect and the model of a femur. Figure 6a shows an ABS FDM model of a skull with a defect while Figure 6b shows one of the fabricated PMMA models to correct the cranial defect placed in the ABS skull. Figures 6a and 6b show PMMA models with different porosity. Figure 7a shows the PMMA model of a femur during fabrication, and Figure 7b shows the final femur model.4. ConclusionsThe building parameters required for successful FDM fabrication with medical-grade PMMA filament (1/16”Ø) were developed using an FDM 3000. It was found that a liquefier and envelope temperature of 235ºC and 55ºC, respectively, as well as increasing the model feed rate by 60%, were necessary to properly and consistently extrude the PMMA filament.

Scaffolds with different porosities and fabrication conditions (tip wipe frequency and layer orientation) were produced, and their compressive mechanical properties were examined. Results show that both the tip wipe frequency (1 wipe every layer or 1 wipe every 10 layers) and layer orientation (transverse or axial with respect to the applied compressive load) used to fabricate the scaffolds, as well as the porosity of the scaffold had an effect on the mechanical properties.Tip wipe frequency had a direct correlation to contraction and inner stress development. Samples fabricated with 1 tip wipe per layer allowed the development of higher inner stresses that resulted in higher yield strength. Although this may be noted as a benefit, one must be cautious of the geometric changes, specifically shrinkage, which may pose problems for models of larger size. The samples fabricated with the high tip frequency had a larger compressive strength and modulus (Compressive strength: 16 ± 0.97 vs.

The samples fabricated with the high tip frequency had a larger compressive strength and modulus (Compressive strength: 16 ± 0.97 vs. 13 ± 0.71 MPa, Modulus: 370 ± 14 vs. 313 ± 29 MPa, for samples fabricated in the transverse orientation with 1 tip wipe per layer or 1 tip wipe per 10 layers, respectively).Specimens fabricated in the transverse orientation yielded desirable mechanical properties due to added support provided by each layer whereas axial orientation specimens only supported loads with every other layer. While porosity did experience a decrease in the transverse orientation compared to the axial orientation, the higher yield strength was seen as a worthwhile benefit. Samples fabricated in the transverse orientation had a larger compressiveabcdFigure 6. Patient-specific PMMA model to repair a cranial defect. a) ABS model of the skull showing the defect. b) PMMA model to correct the cranial defect placed in the ABS skull.

Patient-specific PMMA model to repair a cranial defect. a) ABS model of the skull showing the defect. b) PMMA model to correct the cranial defect placed in the ABS skull. c and d) fabricated PMMA models with different densities.strength and modulus than the ones fabricated in the axial orientation (Compressive strength: 16±0.97 vs. 13±0.83 MPa, Modulus: 370±14 vs. 281±22 MPa, for samples fabricated with 1 tip wipe per layer, in the transverse and axial orientation, respectively). Overall, the compressive strain for the samples fabricated with the four different conditions ranged from 8 – 12%.In regards to the porosity of the samples, in general, the stiffness, yield strength and yield strain decreased when the porosity increased (Compressive strength: 12±0.71 – 7±0.95 MPa, Modulus: 248±10 – 165±16 MPa, Strain: 7±1.5 – 5±1% for samples with a porosity ranging from 55 – 70%). The porosity of the samples was increased by increasing the air gap in the scaffold design.

The porosity of the samples was increased by increasing the air gap in the scaffold design. It was noticed that extremely large air gaps can cause faulty fabrications of highly porous scaffolds. Alternatives to achieving large porosities include using smaller extrusion tips or smaller road widths.The successful FDM fabrication of patient-specific, 3D PMMA models with varying densities, including the model of a structure to repair a cranial defect and the model of a femur, was demonstrated. This work showed the possibility of using FDM for the direct fabrication of PMMA customized structures with varying porosities, and therefore tailored mechanical properties.abFigure 7. Anatomical PMMA model of a femur: a) during fabrication, the highly porous inside can be observed, b) final fabricated model with support.AcknowledgmentsThe research presented here was performed at the University of Texas at El Paso in the W.M. Keck Center for 3D Innovation.

Keck Center for 3D Innovation. Support was provided by the University of Texas System Louis Stokes Alliance for Minority Participation Program under grant NSF-HRD-0703584. The authors are grateful for the assistance of Justin Masias, Eric Anchondo, and Eduardo Basurto of the Keck Center on various aspects of the project. The authors are grateful to Biogeneral, Inc., for participating in the research and providing the PMMA filament used in this study. The authors would also like to thank Stratasys, Inc. for providing the FDM systems used in this study.ReferencesCai, Shengyong, and Juntong Xi. 2008. "A control approach for pore size distribution in the bone scaffold based on the hexahedral mesh refinement." Computer-Aided Design 40, no. 10/11: 1040-1050.Eufinger, H., Wehmoller, M., Harders, A., and Heuser, L. 1995. International Journal of Oraland Maxillofacial Surgery. Vol. 24: 104-110.Gopakumar, S. 2004.

11/12: 1087-1096.Zein, Iwan, Dietmar W. Hutmacher, Kim Cheng Tan, and Swee Hin Teoh. 2002. "Fused deposition modeling of novel scaffold architectures for tissue engineering applications.” Biomaterials 23, no. 4: 1169..

sed Deposition of Ceramics (FDC) and CompositesSeyi Onagoruwa, Susmita Bose and Amit Bandyopadhyay School of Mechanical and Materials Engineering Washington State University Pullman, WA 99164-2920 E-mail: amitband@wsu.eduAbstractFabrication of functional ceramics and composites has been attempted using fused deposition of ceramics process. In this work, first a polypropylene (PP) binder system has been developed for the FDM. The PP binder system was mixed with ceramic powders and then extruded in the form of filament for the FDC. Controlled porosity ceramic parts were directly fabricated using a FDM 1650 machine with mullite, fused silica and titania powder loaded green filaments. The parts were then binder removed and sintered. Some of the porous parts were then infiltrated in Al metal to form the metal-ceramic composites.

The parts were then binder removed and sintered. Some of the porous parts were then infiltrated in Al metal to form the metal-ceramic composites. This article discusses feedstock development, part fabrication and methods to improve the quality of parts during processing.IntroductionRapid prototyping (RP) or solid freeform fabrication (SFF) is a technique in manufacturing in which a solid physical model of a part is made directly from a three-dimensional computer-aided design (CAD) file. There are several RP techniques are available. Some of the most common RP techniques include stereolithography (SLA), selective laser sintering (SLS), fused deposition modeling (FDM), 3-D printing and sanders prototyping. Fused deposition modeling (FDM) is the RP technique used in this study.FDM, a solid-based RP technique, is commercialized by Stratasys Inc. (Eden Prairie, MN).

Fused deposition modeling (FDM) is the RP technique used in this study.FDM, a solid-based RP technique, is commercialized by Stratasys Inc. (Eden Prairie, MN). FDM is an extrusion-based layered manufacturing process in which semi-solid thermoplastic polymers gets deposited on a platform through a nozzle fitted into a heated liquifier controlled in X and Y directions. The cold filament at the top of the liquefier acts as a piston, creating a positive pressure to extrude fine roads of molten material through the nozzle. The motion of the FDM head (X-Y direction) and platform (Z direction) is controlled by the computer, based on the CAD design of the prototype been built. After the head has completed the first layer, the platform indexes down set by the slice thickness and the second layer gets deposited on top of the first layer. This process continues until all layers are complete, leaving a solid model of the desired part.

This process continues until all layers are complete, leaving a solid model of the desired part. There are several variables that controls part quality in FDM. Among the materials related variables critical to FDM include surface chemistry, melt viscosity, stiffness, heat capacity and adhesiveness. Commercially available materials for FDM include wax, ABS and nylon.Fused deposition of ceramics (FDC) is a modified FDM process where ceramic powder loaded thermoplastic filaments are extruded through the FDM [1-4]. During fused deposition of ceramics, thermoplastic polymers melt carry the ceramic powders along with it. The green ceramic parts are similar to injection-molded parts having 40-45 volume % binder in it. The224parts are then subjected to a binder removal and sintering cycle for densification.

The green ceramic parts are similar to injection-molded parts having 40-45 volume % binder in it. The224parts are then subjected to a binder removal and sintering cycle for densification. The process development for FDC work requires optimization of thermoplastic binder composition, filament fabrication, FDC and binder burn out and sintering cycle development.Basic requirements for the development of the FDC feedstock filaments include low viscosity, high strength, high strain and high modulus. The high strength enables the filament to act as a piston during the FDC to prevent buckling. Low viscosity would prevent back flow of molten material. Binder system development is a major key to a successful FDC process. The binder system could be viewed as a parent that nurtures a child until the child is able to live on his/her own.

Binder system development is a major key to a successful FDC process. The binder system could be viewed as a parent that nurtures a child until the child is able to live on his/her own. Good properties of a binder system such as low viscosity, high strength, high strain, high modulus and easy binder burnout would enable the green filament to be easily fabricated in the FDC process into green structures and the green structures to survive the post processing. The properties of ceramic powders introduced into the thermoplastic binder system are very important too. The most important parameters one needs to consider when choosing ceramic powders are surface area, chemical composition, particle size and their distribution. Fine- grained, wide-size distribution ceramic powders lower the overall viscosity of the FDC mixture during compounding.

Fine- grained, wide-size distribution ceramic powders lower the overall viscosity of the FDC mixture during compounding. When developing green filament, it is very desired to have the maximum solids loading of ceramics for reduced problems during post processing.Fabrication of the FDC feedstock such as mullite (3Al2O3, 2SiO2), fused silica (SiO2), titanium dioxide (TiO2), and alumina (Al2O3) involves three distinct processing steps. The first step involves the composition formulation of binder system and compounding of binder system and ceramic powder. The second step involves the extrusion of compounded mixtures into filament. The third step involves the FDC to fabricate high quality, structural parts. Binder removal and sintering are the post operations. Other post operations in the FDC process could include trimming or filing of seam lines in FDC parts while in the green state before binder removal and sintering. Good optimization of material and build parameters (e.g.

Good optimization of material and build parameters (e.g. head speed, flow rate, pre- flow, build temperatures) in the FDM software could prevent some of these post-processing operations.Development of binder system for FDCA binder system was developed from an in-house polypropylene-based FDM feedstock. The binder system was developed from commercially available thermoplastic materials. The prevalent components of the binder system are polypropylene base polymer, wax, tackifier, plasticizer and elastomer.

The prevalent components of the binder system are polypropylene base polymer, wax, tackifier, plasticizer and elastomer. Table 1 summarizes the role of each constituent material in the binder system and the FDC feedstock.Binder constituent/s Base polymer (polypropylene)Characteristics Act as the backbone of the filament and gives strength Enables tackiness and provides flexibility Provides elasticity and flexibility Provides plasticity to filament for spooling Reduces viscosity Table 1: FDC filament constituent/s and their characteristicsTackifier Elastomer Plasticizer Wax225The base polymer gives strength to the entire filament. The role of the base binder is to act as the “backbone” or “strength giver” in the overall FDC feedstock formulation. The base polymer is the most difficult to choose, specifically because the other constituents are chosen to modify the properties of the base polymer. It should exhibit low viscosity, high strength, and good flexibility.

It should exhibit low viscosity, high strength, and good flexibility. The wax component in the FDC filament reduces the viscosity of the filament formulation. The tackifying component or tackifier provides tackiness and also increases the flexibility of the filament composition. The plasticizer provides plasticity and flexibility. It also provides some strength to the filament. The elastomer provides elasticity and flexibility. The elasticity in the filament helps to achieve a uniform feed rate during the FDC operation without any breakage of the filament. The elastomer also enables the filament to be spooled without any significant plastic deformation.Polypropylene, wax, tackifier, plasticizer, and elastomer are general names that represent different groups of materials that behave in a certain way or have almost same characteristics. These materials are produced by different manufacturers with different properties or characteristics.

These materials are produced by different manufacturers with different properties or characteristics. It’s the responsibility of the FDC feedstock developer to hand pick the best brand of polymer and additives that would produce a good binder system and eventual FDC feedstock. Mechanical testing was performed on several brands of materials from different groups and manufacturers. The vendor’s materials with the highest strength, highest strain, and lowest viscosity were selected as the FDC filament constituent for the composition formulation of the binder system.Experiment matrix as a technique of the DOE (Design of experiment) was used to organize and evaluate experimentation to obtain the best binder system composition. The matrix helps better understand what and how each constituent influences the base polymer and the binder system.

The matrix helps better understand what and how each constituent influences the base polymer and the binder system. The matrix was also used to understand how volume percent solids loading of ceramic powders influence the binder system and the corresponding FDC feedstock. The matrix contains results of mechanical testing (strength [MPa], strain [mm/mm], rheology torque [Nm]) of strategic compositions. Results of these various compositions and the outcome of sintered parts fabricated from the FDC feedstock were used to tailor a desirable binder system for the FDC.A binder system has been developed from the above methodology. The binder system has been used to load 55 volume % of mullite, fused silica, and titanium dioxide ceramic powders. The mullite feedstock in volume % comprises of 47.93% mullite powder + 6.85% alumina powder + 0.69% MgO + the developed binder system. The fused silica feedstock comprises of 53% fused silica + 3% MgO + the developed binder system.

The fused silica feedstock comprises of 53% fused silica + 3% MgO + the developed binder system. The titanium dioxide feedstock comprises of 51% titanium dioxide + 4% MgO + the developed binder system. The developed binder system is composed of 44% polypropylene + 13.9% elastomer + 18.7% plasticizer + 7.8% tackifier + 15.6% wax.Filament fabrication for the FDCCompounding of binder system and ceramic powders (mullite, fused silica, and titanium dioxide) was done with the aid of a Haake polylab system (Paramus, NJ). All FDC compositions were compounded at 180oC at roller speed of 120 rpm for 80 minutes in a 48.3 cm3-mixing bowl. The neat binder is first compounded for about 15 to 20 minutes before ceramic powders are added to the binder. Compounded mixtures are allowed to cool down to room temperature before they are226granulated into smaller size particles for extrusion.

Compounded mixtures are allowed to cool down to room temperature before they are226granulated into smaller size particles for extrusion. Ceramic powders especially fused silica tends to absorb moisture from the surroundings, which could reduce the strength of the FDC filament. Granulated compounded mixtures were subjected to moisture removal in a vacuum oven prior to filament extrusion. All the compounded mixtures were extruded at a screw speed of 7 rpm and 170oC (zone1) to 172oC (zone2) to 174oC (zone3) to 168oC (die zone) through a 1.78 mm diameter rod die. A horizontal conveyor belt was attached to the extruder for continuous fabrication of filament and subsequent spooling. The speed of the conveyor belt is very critical in obtaining a filament of 1.78 mm in diameter.Fused deposition of ceramicsThe extruded filaments were used as feedstock for the Stratasys FDM 1650 machine to create high quality parts.

A small length of filament was first used to fill the empty FDM liquefier heated to a particular temperature until the filament is extruded in a continuous semi-molten, string-like manner through the liquefier’s nozzles. These steps are repeated for all temperatures between 200oC and 265oC until a suitable extrusion rate through the nozzles was attained. Set- point temperatures for building ceramic parts from the filaments of mullite, fused silica, and titanium dioxide were established at 235oC and 237oC for both main and support build temperatures, respectively. Envelope temperature of 48oC and material flow rate of 130% was established for part building.FDM machines come with some preset parameters for different materials such as pre-flow, start- flow, main flow, roll back and head speed. These values are preset in Stratasys’ Quickslice software for filaments such as ICW-06 wax and ABS.

These values are preset in Stratasys’ Quickslice software for filaments such as ICW-06 wax and ABS. The FDC filaments have different material characteristics (such as viscosity, melting points, drag coefficient) from the commercially available FDM filaments. New sets of material-dependent values had to be established for the FDC filaments to improve the quality of parts. Using the DOE as a tool coupled with optimization, a new set of material dependent values was established for the FDC filaments (Table 2). Several parts were build using these parameters with mullite, fused silica and titania powder loaded filaments.

Several parts were build using these parameters with mullite, fused silica and titania powder loaded filaments. Figure 1 and 2 show mullite and fused silica parts processed via the FDC.Material dependent parametersPre-flow Start flow Start distance Main flow Roll back Head speedValues for FDC feedstocks: mullite, SiO2 and TiO2 79 37 60 150 153 700Table 2: Material-dependent parameters for the FDC filamentsBinder removal and sinteringBinder removal and sintering are part of the post processing operations. Binder removal and sintering was done in two stages. In the first stage, all the binders were removed. Different227heating rates were used in various temperature intervals to prevent surface and internal cracking of the samples. At low temperatures (< 200oC), liquid binder component(s) were removed via capillary action to the surrounding setter bed of alumina powder in which FDC parts were embedded.

At low temperatures (< 200oC), liquid binder component(s) were removed via capillary action to the surrounding setter bed of alumina powder in which FDC parts were embedded. At a temperature over 200oC, evaporation and decomposition of the residual binders remove the remaining binders. The second stage is the sintering or densification stage. This was done at a temperature of 1650oC and hold time of 3 hours. These two stages of binder removal and sintering helped to fabricate high quality ceramic parts. Linear shrinkage after binder removal and sintering for the FDC part was established at 10-12% for mullite parts and 8-12% for fused silica parts. Figure 3 shows three sintered mullite parts.FDC part qualityVisual inspection and the dimensional accuracy of FDC parts determined the part quality. The dimensional accuracy and tolerance of all the FDC parts were found to be very good in comparison to commercially available FDM feedstock for Stratasys’ FDM 1650 machine.

The dimensional accuracy and tolerance of all the FDC parts were found to be very good in comparison to commercially available FDM feedstock for Stratasys’ FDM 1650 machine. Parts created were very close to the net shape and dimension of the CAD/STL files. All part dimensions were measured within +/- 1% of the original dimensions.Different types of defects in the FDC parts are classified into two main categories: (1) surface defects, (2) internal defects. Surface defects were minimal but visible to the naked eye. Two common surface defects are the staircase and chordal effect. Staircase defect is a common defect of most RP methods and is caused due to layered manufacturing. Reducing the thickness of each layer during operation can reduce this defect. Chordal effect originates in the STL format of the CAD. The STL is very similar to meshing in finite element analysis by approximating the surfaces of the CAD model with triangles. Highly curved surfaces must employ many triangles.

The STL is very similar to meshing in finite element analysis by approximating the surfaces of the CAD model with triangles. Highly curved surfaces must employ many triangles. The smaller the mesh sizes of curved surfaces, the better the quality of the fabricated model. Chordal effect was reduced by specifying binary mode and smallest chord length when exporting CAD files as STL files to the FDM machine. Similar part quality effects have also been discussed in Reference 5.Internal defects need to be addressed seriously if the part is going to be used as a functional component. Internal defects are product of hardware and software specifications and their limitations. The most common types of internal defects are sub-perimeter voids, inter-road voids and core voids. Sub-perimeter voids are created by incomplete filing of areas inside the fused deposition part.

The most common types of internal defects are sub-perimeter voids, inter-road voids and core voids. Sub-perimeter voids are created by incomplete filing of areas inside the fused deposition part. A void is created due to an insufficient material flow to fill the volume at the tangent path and actual perimeter path intersection. Specifying a negative offset (-0.003) to the perimeter in the Quickslice software rectified this defect. Inter-roads voids are created due to the inconsistent material flow of the FDC filaments. This defect is important because of the void it creates in the interior of the ceramic structure. Extruding the FDC filaments to near net size of 1.78 mm diameter, and increasing the flow rate of the FDC filaments to 130% reduced this defect. Core defects were attributed to parts with concentric circular features and created by improper road selection for the contour fill of an area.

Core defects were attributed to parts with concentric circular features and created by improper road selection for the contour fill of an area. Specifying road widths that were ideal to both specified slice thickness and slice interval eliminated these defects. A detailed analysis related to different types of defects and tolerances are presented in Reference 6.228Metal-ceramic compositesOnce the porous ceramic preforms (mullite and fused silica) are fabricated via the FDC process, some of the preforms were infiltrated with molten aluminum metal via pressure-less reactive metal infiltration to form interpenetrating metal-ceramic composite. In these composites, the metal as well as the ceramic phases are connected in all three directions to themselves.

In these composites, the metal as well as the ceramic phases are connected in all three directions to themselves. Some of the advantages of processing the metal-ceramic composites via the FDC route include: (1) microstructure and macro structures of the part can be designed simultaneously in CAD (2) volume fraction as well as phase distribution can be controlled by controlling the pore architecture. Figure 4a and b show the as-processed and polished microstructures of the mullite- Al composites. The uniformity in the metal phase distribution can be seen from these pictures. Both uniform as well as gradient composites can be fabricated by this method.ConclusionsThis study has demonstrated the process of developing a binder system for the FDC process. The developed binder system has been used in fabricating the FDC filaments for mullite, fused silica, and titanium dioxide.

The developed binder system has been used in fabricating the FDC filaments for mullite, fused silica, and titanium dioxide. Green FDC parts were created from the filaments with features and appearance similar to those present in FDM processed wax and polymer parts. The green parts were successfully sintered to obtain fully dense ceramic parts with minimal or no defects. Porous ceramic parts of mullite, and fused silica have been infiltrated with aluminum metals to obtain metal-ceramic composites. Properties of these composite are currently being addressed as a function of their structural parameters.References 1. T. F. McNulty, F. Mohammadi, A. Bandyopadhyay, D. J. Shanefield, S. C. Danforth and A. Safari, “Development of Binder Formulation for Fused Deposition of Ceramics,” Rapid Prototyping Journal, 4 [4], pp 144-50 (1998).2. A. Bandyopadhyay, R. K. Panda, T. F. McNulty, F. Mohammadi, S. C. Danforth and A.

Langrana, A. Safari, P. Whalen, S. C. Danforth, C. Ballard, “Quality of parts processed by fused deposition,” Proceedings. Solid Freeform Fabrication Symposium, Austin, TX, 1996.6. S. Onagoruwa, MS Thesis, Mechanical and Materials Engineering, Washington State University, August 2001.229(a)(b)Figure 1: (a) and (b) Mullite parts fabricated via FDC.Figure 2: Green, brown and sintered fused silica parts via FDC. Three metal infiltrated parts can also be seen in the back of the picture.Figure 3: Sintered mullite parts processed via FDC230(a)(b)Figure 4: (a) As-processed mullite-Al composites. (b) Microstructures of mullite-Al composites.231.

sed Deposition of Ceramics (FDC)for Structural Silicon Nitride ComponentsM.K. Agarwala, A. Bandyopadhyay, R. van Weeren, N.A. Langrana*, A. Safari, and S.C. Danforth Center for Ceramic Research and *Dept. of Mechanical Engineering Rutgers • The State University, Piscataway, NJ.V. R. Jamalabad, P. J. Whalen, R. Donaldson, and J. PollingerAIIiedSignalInc., Morristown, NJ.Fused Deposition of Ceramics is aI! SFF tec~que based on comm~rcial FDMTM ~echno.logy, f?r fabrication of structural and functIonal cerannc components. ThIS study descnbes, ill detail, process improvements made in pre-FDC, FDC, and post-FDC fabrication steps to achieve functional properties . Microstructural characterization of sintered FDC parts reveals microstructures simllar to convennonally processed silicon nitride material. Mechanical properties of FDC processed silicon nitride bend bars and toughness samples were evaluated.

Mechanical properties of FDC processed silicon nitride bend bars and toughness samples were evaluated. These property evaluation studies demonstrate that mechanical properties similar to commercial GS-44 silicon nitride materials can be achieved using FDC. The study also describes results achieved on fabrication of complex components from silicon nitride using FDC.in commercial GS-44 silicon nitride ?omponents.I.IntroductionMost of the SFF techniques have been commercialized for fabrication of polYmer and plastic parts for design verification and form and fit. Recently, some of these techniques are being explored for fabrication of structural metal and ceramic components [1]. Direct fabrication of parts and molds from metals and ceramics using the SFF techniques has created an opportunity to reduce lead times and costs in the development of new products.

Direct fabrication of parts and molds from metals and ceramics using the SFF techniques has created an opportunity to reduce lead times and costs in the development of new products. Such an opportunity for manufacture of new ceramic parts in a matter of days offers several advantages over conventional ceramic processing which can take weeks or months to produce the parts or complex tooling for the parts. The ability to do cost effective, iterative design using SFF techniques enables designers to create components using ceramic design guidelines from the start rather than components designed for metals. Some of these SFF techniques have been developed for ceramic part manufacture and are commercially used in production of parts [1]. One such technique called Fused Deposition of Ceramics (FDC) has been developed and demonstrated for structural ceramics [2-6].

One such technique called Fused Deposition of Ceramics (FDC) has been developed and demonstrated for structural ceramics [2-6]. FDC is based on an existing Fused Deposition Modeling (FDMTM) technology, commercialized by Stratasys Inc., for processing of polymers and waxes [1,7,8].Fused Deposition of Ceramics (FDC)Fused Deposition of Ceramics (FDC) has been developed to create functional ceramic components using ceramic-polymer feedstocks [2-4]. These feedstocks are extruded into filaments of 0.070" nominal diameter, which are then used as the feed material for fabrication of three dimensional green ceramic objects using a commercial FDMTM system, 3D Modeler™. During FDC processing, the polymer/wax acts as a carrier and binder for the ceramic particles as the material flows out of the heated extruder head. The green ceramic object thus created is then subjected to conventional binder removal and sintering processes, to produce fully dense structural ceramic components.

The green ceramic object thus created is then subjected to conventional binder removal and sintering processes, to produce fully dense structural ceramic components. As in the case of any manufacturing process, there are several inter-related process variables which detennine the success and quality of parts fabricated by FDMTM and FDC processes. Poor control and optimization of these variables results in the presence of flaws/voids in the parts which are detrimental to the structural properties of the part.335The FDMI'M process was originally designed for wax and polymer materials to serve the purposes of form and fit and positives for investment casting. Consequently, of chief importance to the end users is the external dimensional accuracy and to some extent, surface finish of the final parts. The presence of internal defects in parts built by FDMTM has not been a priority in the development of the FDMTM process.

The presence of internal defects in parts built by FDMTM has not been a priority in the development of the FDMTM process. The prevention or elimination of these defects, such that ceramic components with functional engineering properties could be made by FDC, is the focus of current FDC research [5,6]. As will be shown here, structural properties have been obtained in a commercial grade silicon nitride through process control at every stage of the FDC process. Further, it is shown that it is possible to rapidly fabricate complex shaped ceramic components.II. Development of Green Filaments for FDCSuccessful FDC processing starts with fabricating filaments of 0.070" ± 0.001" diameter to be used as feed material in the FD hardware. However, for a filament material to be suitable for FD processing, it must possess certain thermal and mechanical properties [4,8].

However, for a filament material to be suitable for FD processing, it must possess certain thermal and mechanical properties [4,8]. The key variables that require careful attention and simultaneous optimization in developing the filaments for FDC processing are: viscosity and adhesion behavior of the material, and the flexural modulus and strength of the filaments. By tailoring the ceramic particle characteristics andlor binder chemistry the properties can be optimized in ways similar to that used by the ceramic injection molding industry. Based on the constraints imposed by these variables for FDC processing, a series of thermoplastic binders have been developed to enable FDC processing. These binders, called the RU series of binders, are four component systems in which the amount of each component was tailored to achieve appropriate viscosity, adhesion behavior, flexibility and stiffness in the final ceramic filament.

Table I shows the four components of the RU series of binder, the role each component plays in the mixed system, and weight percent range of each component. Appropriate tailoring of these components along with the selection of a suitable surfactant for the specific particle system being used have enabled FDC processing of several ceramic systems [3].The work reported here is based on FDC using an in-situ reinforced silicon nitride, referred to as GS-44*. The GS-44 powder is pre-treated with 3 weight % oleyl alcohol which serves as an appropriate dispersant for GS-44 [4]. RU9 and RUl1 binder formulations were used in this study for GS-44. RU9 formulation contains 20.25% wax, 18.75% polymer, 34.5% tackifier, and 26.5% elastomer while the RUll formulation contains 27% wax, 30% polymer, 21 % tackifier, and 22% elastomer by weight. Mixing of the required amount of GS-44 powder, pre-treated with dispersant, with the binders was done using a torque rheometer mixer.

Mixing of the required amount of GS-44 powder, pre-treated with dispersant, with the binders was done using a torque rheometer mixer. GS-44lbinder feedstocks containing 55 volume % GS-44 were used for FDC processing. The compounded feedstock was granulated and sieved for filament extrusion. Continuous lengths of 0.070" ± 0.001" diameter flexible filaments were fabricated using a single screw extruder and wound onto a spool [4].Ill.FDC ProcessingFDC processing, using commercial FDMTM systems, employs the same build strategies used in FDMTM processing of polymers. First, material is deposited as a perimeter defining the boundaries of a given layer of the part followed by raster motion of the liquefier head depositing material inside these boundaries as a series of adjacent roads. This two step strategy is repeated layer-by-Iayer, creating a three-dimensional green ceramic part.

This two step strategy is repeated layer-by-Iayer, creating a three-dimensional green ceramic part. Therefore, any debonding between adjacent roads and adjacent layers, or incomplete filling of any regions of the part will result in property limiting defects in the final sintered ceramic parts. Strength or property limiting defects in FD processed parts have three primary causes.AlliedSignal Ceramic Components, Torrance, California336Surface defects such as, surface finish, stair step formation, etc. due to limitations in the current state of all SFF technologies.Internal defects due to poor optimization of feed filament material, i.e. non-uniform flow rate.Internal defects due to current FDMTM hardware and software limitations.Surface defects arising in FDC parts due to poor surface finish inherent to SFF processing have been eliminated via post-FDC surface finishing operations, as is a common practice in most SFF processing applications.

As shown in Figure 1, the green FDC GS-44 Si3N4 parts fabricated using RU binders do not exhibit any delamination or any debonding between adjacent roads. Several Microstructurally, individual roads and layers are indistinguishable in the green state. internal defects were, however, observed in the early FDC processed parts. Defects arising due to poor control of filament diameter and poor FDC process optimization have been eliminated by better control of filament diameter (to ± 0.001 ") during fabrication by single screw extrusion and by optimization of FDC processing conditions. However, two key defects remained, namely, sub perimeter voids and inter-road defects. These were largely due to limitations in the current FDMTM build strategy and FDMTM system hardware and software.

These were largely due to limitations in the current FDMTM build strategy and FDMTM system hardware and software. Defects arising due to these limitations have also been observed in commercial FDMTM wax and polymer materials [5].Sub-Perimeter Voids: During raster filling of a layer, voids are left due to incomplete filling at points where the raster segment turns around at the perimeter. The FDMTM system generates tool paths with sharp turns in the build file for complete filling at these locations. However, due to the constant velocity of the liquefier, the actual liquefier tool path during deposition is curved. The result of such an action is incomplete material fill at these intersections. The size and occurrence of the sub-perimeter void depends on the build conditions (road width and thickness, and liquefier head speed) and the angle between the raster line and the Perimeter curve [6].

Figure 2 shows schematically how these voids are created and strategies that have been implemented to eliminate them, discussed below.Since sub-Perimeter voids are predictable and the cause of their existence is known exactly, several solutions are feasible for eliminating them. A simple way of eliminating these defects is to give a negative offset to the perimeter. This results in the liquefier head "plowing" into the perimeter during raster motion, as shown in Figure 2b, ensuring void-free parts. However, the negative offset to the perimeter needs to be optimized for each build condition such that the "plowing" action does not result in excessive flow of material at these locations. An alternative solution is to alter the tool path such that the actual tool path does not leave any unfilled regions. This approach includes extending the corners of raster segment turns, at the perimeter, closer to the perimeter [6], Figure 2c.

This approach includes extending the corners of raster segment turns, at the perimeter, closer to the perimeter [6], Figure 2c. Such a change in the tool path in the build file results in an actual tool path motion which is a closer approximation of the sharp tum rather than the curved turn. This solution of altering the path of raster segment at the intersection with the perimeter has recently been developed and implemented successfully for FDC.Inter-Road Defects: Defects between adjacent roads (inter-road defects) can occur in one of two following forms: 1) if there is no physical contact between adjacent roads, defects will be present as voids between adjacent roads, which can be easily observed in the green state, or 2) if there is physical contact between adjacent roads but the bonding between them is not strong, then the defect between adjacent roads may be present as a weak interface.

Such a weak interface is not easily observed in the green state but can cause serious damage to the structural integrity of the part during binder removal and/or sintering. These inter-road defects result from various sources.Poor bonding between adjacent roads can arise from inconsistent material flow due to variation in filament diameter or filament slippage between rollers. This has been overcome by fabrication of filaments with controlled diameter and by optimization of the roller drive system,337Inter-road defects also arise due to excessive cooling of previously deposited respectively. adjacent roads when the road vector length is large. This is analogous to poor sintering between adjacent scan lines in SLS or print lines in 3D Printing when the vector lengths are large [1]. As done in SLS and 3D Printing, bonding between long vector length raster segments in FOC parts has been improved by providing a negative gap between roads such that the roads partially overlap.

As done in SLS and 3D Printing, bonding between long vector length raster segments in FOC parts has been improved by providing a negative gap between roads such that the roads partially overlap. Further improvements in inter-road bonding have been made by optimizing the build environment temperature and optimization of the amount of tackifier in the binder chemistry to provide for optimum adhesion and hence bonding characteristic.Defect Free FDC Processing: A series of square cross-section (1/2" X 1/2") bars were fabricated using a slice thickness of 0.01", road width of 0.02", negative gap between roads of 0.002" and a negative offset was provided to the perimeters to account for the sub-perimeter voids. FOC of GS-44/RU binder systems was done at 150°C-155°C at a liquefier head speed of 0.5 "/second. These bars were then used for mechanical property evaluations after binder removal and sintering.

FOC of GS-44/RU binder systems was done at 150°C-155°C at a liquefier head speed of 0.5 "/second. These bars were then used for mechanical property evaluations after binder removal and sintering. Fabrication of complex components by FOC was demonstrated by fabrication of a top section of a tactical missile radome. The radome .stl file was scaled up by 20% in all three directions to accommodate the shrinkage during sintering and surface finishing operations. The file was then sliced using a slice thickness of 0.01" and then truncated in the Z-direction to obtain only the top 4.8" of the scaled-up version of the radome. The scaled-up radome had a base diameter of 3.1", wall thickness of 0.075" and an angle of inclination of approximately 72°. The green radome was fabricated in 12 hours without the need for any support structure using the same FOC conditions used for 1/2" square bars mentioned above.

The green radome was fabricated in 12 hours without the need for any support structure using the same FOC conditions used for 1/2" square bars mentioned above. However, no negative gap was provided between roads as the vector lengths of raster segments were short enough to allow complete inter-road bonding. In addition to bars and radome, several other simple and complex shaped parts were f~bricated, Figure 3.IV. Binder Removal and Sintering of FDC PartsAs mentioned earlier, 'the RU family of binders being used for FOe are multi-component with different thermal degradation temperature ranges, Table I. For these binder systems, binder removal was done in two stages. In the fIrst stage (room temperature to 450°C), 90 to 95 weight % of the binder was removed in a flowing nitrogen environment. During this stage, different heating rates were used in different temperature intervals to prevent surface as well as internal cracking of the samples.

During this stage, different heating rates were used in different temperature intervals to prevent surface as well as internal cracking of the samples. At low temperatures « 200 °C), liquid binder component(s) were removed via capillary action to the surrounding setter bed material of activated charcoal, in a process called wicking. Such a loss of binder by capillary action creates surface connected internal pore channels that make further removal of binder much easier and prevents sample cracking during the process. At a higher temperature (200°C to 450°C), the remaining binder was removed by evaporation and decomposition. The residual binder (5 to 10 weight %) after the first stage, mostly present in the form of carbon, was removed in the second stage of the binder removal process by placing the sample in an alumina crucible without any setter bed and heating in air to 500 °C. The heating rate and the soaking time at different stages are dependent on the size and shape of the part.

The heating rate and the soaking time at different stages are dependent on the size and shape of the part. The maximum cross-section FOC samples which have been successfully processed for binder removal without any defects are currently limited to 1/2"Xl/2".The GS-44 raw material powder contains primarily a-Si3N4 and a small volume «10%) of oxide sintering aids. During high temperature gas pressure sintering, these oxide additives melt and provide a liquid phase for densification of the porous compact. As densification proceeds, the a-Si3N4 transforms to ~-Si3N4 through a solution-reprecipitation process. When densification is complete, the sintering aids remain in the grain boundaries as amorphous phase~.

When densification is complete, the sintering aids remain in the grain boundaries as amorphous phase~. Sintering of the radome section was done in a machined graphite flXture to prevent any deformatton of the part.338V.Physical and Mechanical PropertiesDensity: The properties of green, brown, and sintered FDe square cross-section bars were compared with similar bars fabricated by extrusion of GS-44 bars with the same binder chemistry used for FDC. Binder removal and sintering procedures for extruded bars were the same as those for FDC parts. Iso-pressed GS-44 samples were used as control samples for property comparison with FDC and extruded samples, since iso-pressing of GS-44 is a conventional manufacturing approach for GS-44 silicon nitride components. Bulk densities of green FDC parts initially were 85%-90% due to internal build defects arising from FDMTM build strategy, discussed earlier.

Bulk densities of green FDC parts initially were 85%-90% due to internal build defects arising from FDMTM build strategy, discussed earlier. After implementation of novel strategies to eliminate the FDMTM related build defects, the bulk densities the sintered of green FDC parts were consistently greater than 96%. As shown in Table II, densities of FDC samples, after implementation of strategies to eliminate build defects, were comparable to those of extruded and iso-pressed samples.Shrinkage: Shrinkage measurements in FDC samples were made to detect anisotropy between the build plane (X-Y plane) and the build direction (Z axis). Significant anisotropic shrinkage behavior is commonly observed in parts fabricated by SFF techniques where the shrinkage is greater in the build direction than in the build plane [1]. In the FDC samples in this study, a similar anisotropic shrinkage behavior was observed between the build direction and the build plane.

In the FDC samples in this study, a similar anisotropic shrinkage behavior was observed between the build direction and the build plane. Before elimination of FDMTM related build defects, the linear sintering shrinkage in the build plane was 12-15% while that in the build direction was 17-20%. However, after implementation of the defect eliminating strategies, this anisotropy in shrinkage has been reduced with average build plane shrinkage of 16.6 ± 1.3% and the build direction shrinkage of 19.3 ± 1.6%. It has been observed that the degree of anisotropy in shrinkage is directly related to the total defect population in the parts. Parts with low defect population exhibit a smaller degree of anisotropy in shrinkage compared to parts with higher defect population, since most of the defects typically exist between layers. Therefore, as the defect population decreases, the parts' greater isotropy in the green state translates to more isotropic shrinkage upon sintering.

Therefore, as the defect population decreases, the parts' greater isotropy in the green state translates to more isotropic shrinkage upon sintering. Sintered samples with densities greater than 98% theoretical density show a difference of 2-3% in shrinkage between the build direction and build plane, which is far less than reported by other SFF techniques, suggesting a more uniform microstructure in FDC processed green parts.Mechanical Property: ASlM MilSpec B 4-point bend tests were conducted on samples machined from extruded, FDC and iso-pressed GS-44 bars. As shown in Table II, average bend strengths of 824 ± 110 MPa measured for FDC samples are comparable to those for extruded and iso-pressed samples. Three point bend test were also carried out with indented samples where an indentation was created by applying a 20 Kg load using a Vickers indentor on the tensile surface of the sample.

Three point bend test were also carried out with indented samples where an indentation was created by applying a 20 Kg load using a Vickers indentor on the tensile surface of the sample. This test gives a qualitative idea of the fracture toughness of the samples where a higher failure strength indicates a higher fracture toughness. As shown in Table II, average indented failure strengths indicate similar fracture toughness behavior for FDC, extruded and iso pressed samples.Microstructure-: Sintered FDC parts were observed under optical and scanning electron microscopes (SEM) to evaluate the microstructures for the following: (a) FDC build related defects, (b) delamination during binder removal and sintering, (c) evidence of road and layerwise build strategies, and (d) GS-44 Si3N4 microstructure. Figure 4 shows SEM fractographs of FDC and iso-pressed samples tested under indented three point bend loading.

Figure 4 shows SEM fractographs of FDC and iso-pressed samples tested under indented three point bend loading. Although some FD defects can be observed on the fracture surfaces of FDC samples, they didn't have a significant effect on the failure strength. Other than a few FD related defects, the fracture surfaces of conventionally and FDC processed samples appear similar. The microstructures in the directions perpendicular and parallel to build direction were similar. Individual roads and layers were indistinguishable. There were no discernible signs of layerwise fabrication.339As shown in Figure 5, the sintered FDC radome section maintained its overall conical shape and straight sides with a sintered angle of inclination of 71 0 • The sintered wall thickness was 0.065" accounting for 13.30/0 linear shrinkage in X-Y build plane. The sintered base diameter was 2.6" leading to measured shrinkage in the diameter at base of 16.1 %.

The sintered base diameter was 2.6" leading to measured shrinkage in the diameter at base of 16.1 %. The shrinkage in build direction (Z) of 20.8% was slightly greater than expected shrinkage of 20%, with the fmal vertical height of 3.8". Although, the density of the radome was measured to be greater than 98% of theoretical density of OS-44, several defects were observed on the surface and internally, detected using X-ray radiography. Most of these defects were primarily due to the FD build approach used.VI.ConclusionsoS-44 silicon nitride parts have been formed using a new method, Fused Deposition of Ceramics, FDC. Simple and complex shapes have been produced with high green density after most of the FDC build defects were eliminated. The process has been developed to produce green parts with largely defect-free internal microstructures.

The process has been developed to produce green parts with largely defect-free internal microstructures. It has been demonstrated that FDC green parts can have the binder removed and be sintered to densities and strengths comparable to those produced by conventional ceramic processing approaches. Development of such a process as FDC will enable ceramic design engineers to rapidly fabricate complex shapes and verify the design for not only form and fit but also for performance.Acknowledgmen tsThis work has been done under the support ofDARPA/ONR contract No. N00014-94-0115. The authors would like to thank: Dr. Clifford P. Ballard of AlliedSignal Research and Technology; Mr. William R. Priedeman and Jim Comb of Stratasys Inc., and Mr. Ron Knight and Joe Wright of Lockheed Martin Inc., for their support and help in this work.References1. Proceedings of the Solid Freeform Fabrication Symposium, Vol. 1990, 1991, 1992,1993, 1994, 1995, Edited by H.L. Marcus, J.J. Beaman, J.W.

wFigure 4: SEM fractographs of indented 3-point bend samples fabricated by (a) FDC (arrow indicates the build direction) and (b) iso-pressing.Figure 5: Sintered FDC OS-44 samples, including two tactical missile radome sections.344.

The particulate systems explored vary in particle size from nano-crystalline (WC-Co) to coarse (>100 Jlm Si(2) particles. The material systems explored for FDC and FDMet vary from conventional ceramic and metal systems such as Si02 and stainless steel to advanced materials such as Si3N4 and PZT. FDC and FDMet of such a variety of material systems using a commercial FDMTM system has been made feasible by development of a unique series of binders, as well as optimized FD processing, binder removal techniques and sintering conditions.I.IntroductionSFF techniques have been commercialized for fabrication of polymer and wax parts for design verification and form and fit [1]. Wax and polymer parts made by SFF have also been used to produce metal and ceramic parts using a two-step process. Once a wax or polymer part is built by SFF, it serves as the positive for investment casting of metallic alloys [1].

Once a wax or polymer part is built by SFF, it serves as the positive for investment casting of metallic alloys [1]. Also the wax and polymer materials are being used to make molds for ceramic slurry and gel casting [2,3]. The direct production of metallic or ceramic prototypes for functional applications and testing is currently a major focus in the field of SFF [1,4-7].Most of the SFF techniques used for fabrication of metal and ceramic parts employ polymeric binder systems to bond the ceramic or metal particles together to form a green part [1,3 7]. Once a green part is formed by the SFF technique, further processing follows the approach of conventional green forming techniques, such as powder injection molding. The fabricated green part is processed to remove the binder from the part and then sintered or infiltrated with a lower melting second phase.

The fabricated green part is processed to remove the binder from the part and then sintered or infiltrated with a lower melting second phase. However, fabrication of green metal or ceramic parts by various SFF In the SFF techniques, effort is required in developing the techniques have been far from trivial. proper binder chemistry and often tailoring the as-received commercial ceramic or metal powders to allow SFF processing. Once the binder and particulate systems have been developed, the SFF processing conditions are optimized to fabricate high quality green bodies.This article discusses the results of a new SFF technique, called Fused Deposition of Ceramics (FDC) and Metals (FDMet), which has been applied to a wide range of ceramic and metal systems [3,6,7]. FDC and FDMet are based on an existing Fused Deposition Modeling (FDMTM) technology, commercialized by Stratasys Inc.

FDC and FDMet are based on an existing Fused Deposition Modeling (FDMTM) technology, commercialized by Stratasys Inc. for polymers and waxes [8,9].II.Fused Deposition of Ceramics (FDC) and Metals (FDMet)Fused Deposition of Ceramics (FDC) and Metals (FDMet) is being developed to create functional ceramic and metal components using ceramic or metal powders mixed with organic binder systems. The mixed powder-binder feedstock is extruded into filaments of 0.070" nominal diameter, which are then used as the feed material for fabrication of green ceramic or metal parts385using a commercial FDMTM system. The FDC/FDMet green part is then subjected to conventional binder structural ceramic/metal components.

The FDC/FDMet green part is then subjected to conventional binder structural ceramic/metal components. As in the case of any manufacturing process, there are several inter-related process variables which determine the success and quality of parts fabricated by FDC or FDMet.removal and sintering processesto produce fully denseSuccessful FDC/FDMet processing starts with fabrication of filaments of uniform diameter (0.070" ± 0.001") to be used as feed material in the FD hardware. However, for a filament material to be suitable for FD processing, it must possess certain thermal and mechanical properties [9,10]. The key variables that require careful attention and simultaneous optimization in developing the filaments for FDC/FDMet processing are: viscosity and adhesion behavior of the material, and the flexural modulus and strength of the filaments.

Based on the constraints imposed by these variables, a series of thermoplastic binders have been developed to enable FDC/FDMet processing. These binders, called the RU binder series, are four component systems - an elastomer, tackifier, wax, and a polymer. The amount of each component was tailored to achieve appropriate viscosity, adhesion behavior, flexibility and stiffness in the green filaments [10]. Appropriate tailoring of these components and selection of a suitable surfactant for a specific particle system have enabled FDC/FDMet processing of several ceramic and metal systems.FDC/FDMet using the RU series of binder has been demonstrated for Si3N4, fused Si02, A1203, In each of these particle lead zirconium titanate (PZT), stainless steel and WC-Co. systems, initial trials for FD were done using a basic composition of the RU binder series, called RUl, which has 30% wax, 35% polymer, 20% elastomer, and 15% tackifier by weight.

systems, initial trials for FD were done using a basic composition of the RU binder series, called RUl, which has 30% wax, 35% polymer, 20% elastomer, and 15% tackifier by weight. Further optimization of the binder from the RUI composition was done, if needed, to achieve the needed flexibility in the filaments to allow automated FD processing or to lower the viscosity to enable FDC/FDMet processing. As reported elsewhere [10], the binder chemistry has been nearly optimized for Si3N4 to Yield flexible and continuous filaments. FDC/FDMet of other materials has been demonstrated without much variation in the binder composition from the RUI chemistry. Mixing of the ceramic or metal powder with the binders was done in a torque rheometer mixer to achieve 50 to 65 volume % particle loading. The particle loading achieved in a specific particulate system depends on the particle size and particle size distribution and whether or not a suitable surfactant is used.

The particle loading achieved in a specific particulate system depends on the particle size and particle size distribution and whether or not a suitable surfactant is used. These factors determine the viscosity of the mixed feedstock which should not exceed a certain limit which will prevent FD processing. Filaments were fabricated from the mixed feedstock using a capillary rheometer or a single screw extruder [10].Once suitable filaments were fabricated, FDC/FDMet processing was done using a commercial FDMTM system, 3D Modeler™, hardware and software. Build strategies used in the FDC/FDMet were largely the same as those for FDMTM processing of wax and polymers. However, some novel build strategies were developed and implemented to prevent and eliminate property limiting internal defects which arise in conventionally processed FD parts [11].

However, some novel build strategies were developed and implemented to prevent and eliminate property limiting internal defects which arise in conventionally processed FD parts [11]. The choice of slice thickness and road widths for FDC/FDMet were determined by the nozzle diameter, which in turn is determined by the maximum particle size in the specific ceramic or metal system. The FDC/FDMet green parts were further processed to remove the RU binder completely. The binder removal process is done in two stages such that the RU binder is completely removed without any cracking or damage to the part [6,7].

The binder removal process is done in two stages such that the RU binder is completely removed without any cracking or damage to the part [6,7]. The sintering of the FDC/FDMet part is done by conventional sintering operations for the specific particle system used under consideration.FDC of Silicon NitrideAlthough several ceramic and metal systems have been investigated for the FDC/FDMet processes, the most extensive work to date has been done using an in-situ reinforced Si3N4 from AlliedSignal Ceramic Components of Torrance, California. As-received Si3N4 powder coated with appropriate surfactant was used in this study with 55 volume % Si3N4 in the mixed powderlbinder feedstock. The RU series of binder used for FDC of Si3N4 has been optimized to386result in flexible and continuous 0.070" ± 0.001" diameter filaments fabricated by single screw extrusion [10].

The RU series of binder used for FDC of Si3N4 has been optimized to386result in flexible and continuous 0.070" ± 0.001" diameter filaments fabricated by single screw extrusion [10]. FDC of Si3N4 has been done successfully using novel build strategies employed in sintered parts with physical and mechanical with a commercial FDMTM system to result properties comparable to those obtained by conventional processing of Si3N4 [7]. FDC of Si3N4 has also been demonstrated by fabrication of complex engineering components. Details of the FDC process for Si3N4 and their results can be found elsewhere [7].FDC of Fused SilicaFused silica, alumina, and zircon are commonly used for fabrication of investment casting tooling. Traditionally, investment casting ceramic shells and cores are manufactured using machined positive patterns of the parts or by injection molding of the actual cores and shells.

Traditionally, investment casting ceramic shells and cores are manufactured using machined positive patterns of the parts or by injection molding of the actual cores and shells. SFF techniques have been used directly or indirectly for fabrication of investment casting cores and shells. Indirect fabrication involves SFF fabrication of the positive pattern from wax and polymers for the mold. The wax or polymer positive is then used to form a ceramic shell by sequential dipping of the pattern in a ceramic slurry [1]. The wax or polYmer is then melted out of the shell followed by firing the shell. Direct fabrication of such ceramic shells and cores for investment casting has been commercialized using the 3D Printing process [5].

Direct fabrication of such ceramic shells and cores for investment casting has been commercialized using the 3D Printing process [5]. The study reported here demonstrates the feasibility of using the FDC process for fabrication of silica parts for investment casting with properties comparable to parts fabricated by conventional core making techniques.A commercial grade silica powder used for injection molding of investment casting cores by Certech Inc. of Woodridge, New Jersey, was used for FDC in this study. The particle size of the as-received silica powder was in the range of O.OIJ.lm - 150J.lm, with an average particle size of 70 J.lm. Conventionally, injection molding of cores done by Certech uses 68 volume % silica such that linear shrinkage during fIring is limited to -1 %, with a final density of -70% of theoretical density. In order to have comparable shrinkage and final density in the FDC parts, the RUI binder was mixed with 65 volume % silica powder.

In order to have comparable shrinkage and final density in the FDC parts, the RUI binder was mixed with 65 volume % silica powder. Filaments fabricated by capillary and single screw extrusion were straight and stiff with insufficient flexibility for continuous winding onto a spool. A series of 0.25" X 0.25" X 2.5" bars for modulus of rupture (MOR) tests and simple core shape were fabricated by FDC, Figure 1. After binder removal, the bars were frred at Certech Inc. at 1225°C for 6 hours, Figure 2. MOR testing was done to evaluate the mechanical properties. The bars were also used for physical property evaluations, which are critical in determining whether or not the parts can be used for investment casting applications.

The bars were also used for physical property evaluations, which are critical in determining whether or not the parts can be used for investment casting applications. These physical and mechanical properties were compared with those of injection molded samples of the same grade of silica.As shown in Figures 3, the microstructure of frred FDC silica samples is comparable with that of frred injection molded samples with no evidence of delamination or inter-road debonding. Similarly, the physical and mechanical properties of the FDC processed samples compare very favorably with the reported properties of injection molded parts, Table I, and were within the commercial upper and lower acceptable limits reported by Certech's specification data sheet. For example, although the average cristobalite level in FDC samples is lower than those in injection molded samples, the levels are well within the acceptable limits of 20% and 7%, respectively.

For example, although the average cristobalite level in FDC samples is lower than those in injection molded samples, the levels are well within the acceptable limits of 20% and 7%, respectively. The starting volume fraction of silica in FDC samples is lower (65%) than that in injection molded parts (68%), while the final bulk density of both FDC and injection molded parts are similar. Therefore, the linear shrinkage observed in FDC samples was slightly higher. Since the number of FDC samples tested was few (10) in comparison to injection molded samples used for reporting data in product data sheet, the standard deviation for the properties in FDC samples (shown in parenthesis in Table I) indicate a wider spread in the FDC values compared to the injection molded samples.

However, this study demonstrates that FDC is a viable SFF technique for direct fabrication of investment casting ceramic cores and shells using commercial grade powders with resulting properties comparable to those of conventionally processed parts.387FDC of Piezoelectric ceramicsPiezoelectric materials have the ability to convert electrical energy into mechanical energy or conversely convert mechanical energy into electrical energy. The applications for devices from these materials speakers, accelerometers, strain gages, ignitors, and various medical diagnostic systems. Several ceramics exhibit piezoelectric behavior, with lead zirconium titanate (PZT) being the most extensively used. Piezoelectric ceramic/polYmer composites, consisting of a piezoelectric ceramic, such as PZT, in an inactive polymer, have shown superior properties when compared to monolithic single phase materials.

The piezoelectric properties have been vastly improved over the last two decades through fabrication of numerous innovative structures [12]. However, most of the processing techniques lack the flexibility of fabricating novel structures.includetransducers, microphones,phonographicpick-ups,SFF techniques, specifically FDC in this study, provide an opportunity to control the fabrication of these complex structures as material is deposited one volume element at a time only in desired locations and is directly controlled by the computer through the component CAD file [1]. Such flexibility, of SFF in general and FDC in particular, allows deposition of material at a very fme scale only in desirable locations and leaving the other areas unfilled [3]. Once the desirable green ceramic structure of PZT is fabricated, the remaining unfilled regions are filled by the inactive polymer phase, following binder removal and sintering.

Once the desirable green ceramic structure of PZT is fabricated, the remaining unfilled regions are filled by the inactive polymer phase, following binder removal and sintering. Such an approach results in complex geometries of piezocomposites to be fabricated. with relative ease. In this study, as received spray-dried PZT powder from Morgan Matroc Inc. of Cleveland, Ohio, was used. Green PZT filaments containing 50 volume % PZT powder and 50 volume % RU1 binder with a surfactant and plasticizer was used, Figure 4. Several simple shapes with intricate internal structures, such as ladder structures, were formed using FDC, Figures 4 and 5. After binder removal and sintering of the green FDC PZT parts, the parts were infiltrated with an inactive the polYmer to form the fmal desired piezocomposites. piezocomposites by FDC were measured and found to be significantly better than those reported for conventionally processed PZT composites.

piezocomposites by FDC were measured and found to be significantly better than those reported for conventionally processed PZT composites. Further details and results on FDC of piezoceramics can be found in these proceedings [3].The piezoelectric properties ofFDMet of Stainless SteelAs mentioned earlier, one of the major applications of SFF technologies has been to fabricate polymer or wax positive patterns for investment casting of metallic alloys. Such an approach has been successfully adopted by automotive and aerospace industries in investment casting of critical components from metals [1]. To further maximize the benefits of SFF technologies, direct fabrication of such metal components by SFF techniques are being developed.

To further maximize the benefits of SFF technologies, direct fabrication of such metal components by SFF techniques are being developed. This study demonstrates the feasibility of using FDMet for fabrication of green parts from powder stainless steel which can then be further processed by binder removal and sintering to result in a fully dense steel part.Precipitation hardened stainless steels are commonly used in various industrial and military applications where resistance to corrosion and high mechanical performance at temperatures to 400°C are necessary. Typically these steels are available in wrought form and the parts are fabricated by conventional thermomechanical processes. The precipitation hardened steel used in this study was a 17-4PH grade (also referred as Type 630) stainless steel powder of -325 mesh particle size.

The precipitation hardened steel used in this study was a 17-4PH grade (also referred as Type 630) stainless steel powder of -325 mesh particle size. Powder with 60% by volume was mixed with RU1 binder composition and extruded into stiff, straight filaments of 0.070" nominal diameter using a capillary extrusion process. Simple green stainless steel shapes were created at 110°C using the commercial FDMTM system, Figure 6. Following binder removal, the parts were sintered at 1350°C for 1 hour in a mild reducing atmosphere (H2+N2), Figure 7. Although sintered parts exhibit high density (92%-95% of theoretical), close examination of sintered parts reveal internal delamination and cracking in the388Further experiments with capillary extruded green rods (3/8" diameter) of the same parts. It is composition revealed cracking and bloating occurring during the binder removal process.

It is composition revealed cracking and bloating occurring during the binder removal process. expected that once the binder removal is optimized to prevent defects, stainless steel FDl\;1et parts with properties and microstructures comparable to that of wrought formed parts can be achieved.FDC of WC-CoTuno-sten carbide (WC)-based composites have a good combination of properties including high hardne~s, toughness, and wear and abrasion resist~nce. ~ese properti~s. make. them suitable for a variety of tooling applications such as metal cutting, mInIng, rock drilling, dies, and wear parts. The basic tungsten carbide:cobalt (WC-Co) 1I!ate~al has been ~odified over the year~ to produce a variety of cemented carbides for these apphcanons. Convennonal powder metallur~cal techniques are used for processing and fabrication for these materials. Today, new techniques are being developed to reduce costs and improve performance.

Convennonal powder metallur~cal techniques are used for processing and fabrication for these materials. Today, new techniques are being developed to reduce costs and improve performance. One of these new technologies include synthesis and use ofnanocrystalline WC-Co powders. As SFF techniques are being explored for metals and ceramics and the SFF technologies have also led to significant growth and demand in Rapid Tooling applications, it is only natural to explore SFF techniques for WC-Co.In this study, FDC was explored as an SFF technique for fabrication of WC-Co shapes. Two different grades of commercial WC-Co powders were used in this study, nanocrystalline WC-15Co from Nanodyne, Inc. and WC-15Co from Valenite, Inc. Due to the extremely fine particle size and high density (-14 g/cm3) of these materials, the viscosity of the RUI binder mixed with only 50 volume% of these WC-Co was too high to allow successful FDC processing using commercial FDMTM systems.

To allow successful FDC processing, the RU series binder was tailored and a suitable dispersant was developed to lower the viscosity of 50-55 volume % WC-Co loaded binder system. Initial FDC trials with the RU binder formulation and use of dispersant, indicate that it is possible to achieve high solids loading (50-55 volume%) of nanocrystalline WC Further Co and attain a viscosity level development of binder and dispersant are expected to result in optimized FDC processing of WC Co with sintered properties comparable to those of conventionally processed WC-Co.low enough for FDC processing at 200°C [10].FDC of AluminaAlumina is a commonly used ceramic in many structural applications. FDC of alumina has been demonstrated by Lone Peak Engineering of Draper, Utah, with very much the same approach In the study by Lone Peak Engineering, FDC was done using commercial as reported here [4]. grade A-16SG alumina powders and thermoplastic binders with suitable plasticizers and dispersants.

grade A-16SG alumina powders and thermoplastic binders with suitable plasticizers and dispersants. As discussed elsewhere [4], FDC processed alumina parts in the study exhibited greater than 97% theoretical density after sintering.ConclusionsThis article demonstrates that FDC and FDMet are SFF techniques which are practical for rapid fab?cation ?f structural and functional parts/components fron:: powderlbinder mixtures. A unique senes of bmders developed for FDC/FDMet have been unIformly applied to a wide range of ceramic and metal systems for FD processing. The materials studied here demonstrate that as received commercial powders can be readily used for FDC/FDMet with sintered parts exhibitino" p~op~rties compara~le to th,?se of conventionally process~d parts from the same material.

By tailonng the RU senes of binder and proper selecnon of dispersants and optimization of the FD process and binder removal procedures, different material systems can be readily developed for FDC/FDMet processing.389AcknowledgmentsThe work on various material systems has been done through support from DARPA, ONR (Contract No. 00014-94-0115), and the New Jersey Commission on Science and Technology (NJCST). The authors would also like to thank Drs. C. Ballard and S. Das of AlliedSignal Research and Technology, Drs. E. Krug and G. Carrasquillo of Certech Inc., Mr. W. Priedeman of Stratasys Inc., Drs. R. Sadangi and L. McCandlish of Nanodyne Inc., and Dr. S. Raghunathan of Valenite Inc. for their help and support in developing FDC/FDMet of thse materials.References1.. Proceedings of the Solid Freeform Fabrication Symposium, Vols. 1-7, 1990 - 1996, Edited byH.L. Marcus, J.J. Beaman, J.W. Barlow, D.L. Bourell, and R.H.

sed Deposition of Ceramics: Progress Towards a Robust and Controlled Process for CommercializationR. Clancy, V. Jamalabad, P. Whalen AlliedSignal, Inc. Research and Technology Morristown, NJ 07962P. Bhargava, C. Dai, S. Rangarajan, S. Wu, S. Danforth, N. Langrana, A. Safari Center for Ceramic Research Rutgers - The State University ofNew Jersey Piscataway, NJ 08855The feasibility of using the Fused Deposition of Ceramics (FDC) process to rapidly fabricate functional quality advanced ceramic components has been demonstratedl This direct manufacturing technique, by eliminating the need for costly tooling, dramatically reduces functional prototype development time. This makes it suitable for small quantity production runs and complex parts. The move from 'feasibility" to a robust, reliable commercial fabrication tool requires that every aspect of the manufacturing be understood and brought under control.

The move from 'feasibility" to a robust, reliable commercial fabrication tool requires that every aspect of the manufacturing be understood and brought under control. An overview of the five basic process steps in FDC: batch compounding, filament fabrication, fused deposition, binder burnout and sintering will be presented in light of this drive toward a robust process. Tools such as Statistical Process Control and Experimental Design techniques are being used to monitor, improve, and stabilize each step and sub-process. Hardware and software modifications have been made to the FD machine to effect the required changes. This paper will identify the remaining technical barriers to commercialization and our progress in addressing these issues.5.I.Introductionfabrication of ceramic components using SFF techniques has created an opportunity to reduce lead times and costs in the development of new products.

The Fused Deposition of Ceramics process uses a ceramic/polymer filament as feed material to build a part from a CAD file by additive layering. The binder is then removed and the part is sintered to full density. In order to apply this process to existing commercial manufacturing needs, effort has been directed to reduce variability and increase robustness. A reproducible, predictable process creating final parts of specified dimensions and strength within 6cr deviation is crucial for commercialization.DirectDue to the sequential, "building block" nature of the FDC process, upstream material- or process-variability can result in major part defects such as voids, cracks, and warpage; this is a common issue for conventional ceramic processing. Batch-to-batch variability results in unpredictable filament quality, and operator responses to this deviation often introduce new process variables.

Batch-to-batch variability results in unpredictable filament quality, and operator responses to this deviation often introduce new process variables. It is difficult to isolate the root causes of technical hurdles, such as material contamination, filament buckling, voids and surface irregularities in the FDC parts, and irregular185shrinkage, without establishing a baseline process. However, due to time constraints of the program, many of these issues were attacked concurrently. Where the problems could not be they were quantified, possible root causes were identified, and eliminated or contained, experiments were planned.The fIrst step to process improvement was run-charting, tracking experimental data through time. This evolving database will yield Quality Control measures such as target and limits, introducing Statistical Process Control. Standard Operating Procedures were established to lessen operator-to-operator variability, and environmental controls were digitized.

Standard Operating Procedures were established to lessen operator-to-operator variability, and environmental controls were digitized. The FDC steps were characterized by measuring viscosity, modulus, and material content in each batch. The application of Design of Experiments (DoE) methodology to FDC sub-processes identified effective parameters and optimized conditions quickly.II. Batch Compounding and Filament Fabrication2,3The ceramic system employed in this study is an in situ-reinforced Si3N4, commercially supplied by AlliedSignal (Ceramic Components, Torrance, California) as OS-44. The powder is coated with a surfactant, oleyl alcohol, in a batch ball-milling process. The coated powder is compounded with the proprietary binder system, designated RU9 (Stratasys, Inc., Eden Prairie, Minnesota), in a torque rheometer. The compounded material is granulated and sieved, then fed into a single screw extruder with an attached torque rheometer.

The compounded material is granulated and sieved, then fed into a single screw extruder with an attached torque rheometer. Continuous filament of nominal 0.070" +/- 0.001" diameter is extruded through a die and wound on spools for storage and use in the Fused Deposition process.to ensuring a reproducible FDC process. As such, a number of process characterization techniques monitor the uniformity of the filament fabrication operation. Loss on Ignition (LOI) tests measure the amount of oleyl alcohol (after milling) and binder (after compounding). Torque values are monitored during compounding and extrusion. An Instron capillary rheometer measures viscosity and load by extrusion of the compounded material. The diameter of the filament is monitored by a laser micrometer as it is extruded.A homogeneous mix of powder, surfactant, and binderis crucialThe solvent drying step after ball milling causes a non-uniform distribution of oleyl alcohol, resulting in batch-to-batch variability.

Mixing and sieving of six coated batches together increased homogeneity of the feedstock, as seen in Figure 1. LOI tests after mixing and subsequent compounding show reduced variability due to the mixing and sieving step. In addition, particle size analysis shows that the mixing and sieving step decreased agglomeration, from 50% (coated powder) to greater than 95% particles finer than 1 /lm. This de agglomeration reduced the torque variation during compounding and pressure variation in the capillary rheometer.In order to increase the homogeneity of the filaments, a breaker plate and screen were inserted between the extrusion barrel and die in the single screw extruder. The plate and screen186reduce agglomeration and inclusions in the mix, and increase the pressure at the die entrance, leading to better mixing efficiency during filament fabrication.coatedcompoundedcoated powders mixed and sievedLOI (0.,4 wt.loss) ave.

compoundingof powders 2.93 ± 1.08 2.86 + 0.18torque (m-g) 1071 ± 247 1033 + 52Instron load variation (N) 61 23LOI (% wt.loss) of RU955 19.30 ± 0.33 19.31 ± 0.12Figure 1. Increased homogeneity due to mixing and sieving step.III. Fused Deposition1,3,4is continuously fed between a pair of motorized rollers into a heated liquifier, where the binder melts. The filament acts as both feedstock and piston, pushing the melt through the nozzle and onto the build platform. Software controls the translational motion of the liquifier/nozzle assembly and the height of the platform, thus "writing" the layered part out of extruded "roads". The extruded road hardens as the binder cools. Material is initially deposited as a perimeter, defining the boundary of each layer of the part, then as a fill pattern, either contours or rasters.

The extruded road hardens as the binder cools. Material is initially deposited as a perimeter, defining the boundary of each layer of the part, then as a fill pattern, either contours or rasters. The enclosed build envelope is maintained at a warm temperature in order to promote bonding between adjacent roads.In the process of fused deposition, filamentOne of the most difficult tasks of this program has been the balance of filament properties required by the process: flexible (for spooling), yet stiff (as a piston), low viscosity (reduce shear in the liquifier) but high modulus. A major obstacle to FDC automation has been buckling of the filament under the rollers and above the liquifier. This is caused by low filament modulus, and is exacerbated by high viscosity reducing flow through the liquifier and heat transfer softening the binder.

This is caused by low filament modulus, and is exacerbated by high viscosity reducing flow through the liquifier and heat transfer softening the binder. Direct cooling supplied by an air conditioner onto the filament at the rollers has raised the filament modulus in the necessary process location, and increased the robustness of the filament such that uninterrupted builds of 16 hours have been achieved. The cool air is recirculated in tubes, isolating the cooling process from the warm build envelope; this enables the operator to maintain independent warm and cool environments where they are required.The nozzles supplied by Stratasys have an internal 0.070" channel, which tapers at the bottom to the particular nozzle tip diameter (0.015", 0.025") at a 1180 angle. The liquifier flow at the liquifier-nozzle channel is 0.073" diameter, thus creating two possible dead zones: interface and at the nozzle channel-taper interface.

The liquifier flow at the liquifier-nozzle channel is 0.073" diameter, thus creating two possible dead zones: interface and at the nozzle channel-taper interface. The nozzles were redesigned (Figure 2) with a constant channel taper from 0.073" at the top to the tip diameter at the bottom, with the flow channel angle optimized (A. Yardimci, the University of Chicago) to maximize the flow rate for the specific tip diameters. Application of the machined nozzles lowered the motor torque driving the filament as well as the torque variability, and helped to eliminate nozzle clogging.187Standard NozzleRe-engineered NozzleFigure 2. Nozzles re-designed with optimized internal angle and "dead zones" removed.In order to accurately predict final part dimensions and reduce the necessity for machining, the build roads must be of consistent width and have no apparent seams or bulges at the joints.

For contour and raster fills, inconsistent road widths and poor seams can cause internal voids which limit part strength. For perimeter roads, these issues cause sub-perimeter voids and poor surface quality. The parameters which affect flow in the start- and end- zones of the tool path were identified, and the process was optimized for constant width perimeters and controllable seams, as in Figure 3.Figure 3. DoE identified critical parameters and optimized perimeter consistency.188Two locations of internal part voids l,4 are inter-lamellar (between adjacent roads) and sub-perimeter (caused by deviation from the set toolpath, due to control limitations), Figure 4. Inter-lamellar voids have been addressed by offsetting the road location such that roads overlap by ~20% of their width.

Inter-lamellar voids have been addressed by offsetting the road location such that roads overlap by ~20% of their width. Sub-perimeter voids are also addressed by over-filling, either by raster offset (impinging the raster into a contour or perimeter by ~10% of the latter width) or toolpath alteration to fill the gaps at the corners of the raster turn ("dog-earing"). Figure 4 illustrates the results of experiments to optimize the internal fill of parts, ensuring full green density and consistent strengths.sub-perimeter void at raster/perimeter juctioninter-lamellar void between adjacent rastersFigure 4. Internal voids reduced by flow control, off-setting, and toolpath alteration;(a.) standard build, (b.) build with inter-lamellar correction and dog-earing.IV. Binder Burn-OutThe binder matrix must be removed from the green FDC parts before the ceramic particles can be sintered to full density.

Binder Burn-OutThe binder matrix must be removed from the green FDC parts before the ceramic particles can be sintered to full density. Thermal degradation occurs during Binder Burn-Out (BBO), a two stage process in which most of the binder is removed slowly at moderate temperatures in a N2 atmosphere, followed by rapid degradation of residual carbon at higher temperature in air. The parts may crack if binder is removed too quickly, and reproducible shrinkage is critical for meeting final part dimensional specifications.In order to characterize the critical temperatures in the binder removal process, draw trials were used to measure weight loss and shrinkage of samples soaked at incremental temperatures. The results in Figure 5 show maximum shrinkage at low temperatures with minimal binder weight loss. This may be due to microstructural rearrangement toward ceramic particle-particle less shrinkage occurs as more binder contact and void removal.

This may be due to microstructural rearrangement toward ceramic particle-particle less shrinkage occurs as more binder contact and void removal. At higher temperatures,189volatilizes, creating residual pores. Draw trial results were incorporated into the BBO schedule (170 0 C dwell to maximize shrinkage) and a DoE (critical intermediate temperatures as binder degrades) which considered the effects of ramp rates, environment, and sample size on shrinkage and cracking. Results indicate that extended dwell times increase shrinkage and lower ramp rates eliminate cracks. The BBO schedule was optimized to ensure reproducible, predictable shrinkage with no part cracking.1.8[I standard de\ • average (x,Y z)ation10090 80~ C> ..:l ........ ..:l ~0I I0 trl ......I J0 0 C'lI + 4- +I0 trl C'l0 0 ("')0 trl ("')0 ~0 trl ~~ 70 ~ 10 ~ 60 ..c 50 .... b1l 40 ~ ~ 30 ?J. 20 1000100200300400Temperature (C)Temperature (C)Figure 5.

b1l 40 ~ ~ 30 ?J. 20 1000100200300400Temperature (C)Temperature (C)Figure 5. Draw trials of Binder Burn-Out process show critical temperatures.The application of process improvement tools such as quality control and DoEs, from feedstock and filament fabrication through fused deposition and binder removal, have reduced shrinkage variability to ± 0.5%. This controlled, reproducible shrinkage from FDC to final sintered component (Figure 6) allows a more accurate CAD file estimation and reduces the need for machining of the part, which increases cost and can introduce damage.V. Mechanical Properties of FDC SamplesIn order to manufacture functional ceramic parts, strengths comparable to conventional ceramic processing are required. Also, the strengths must be reliable and predictable, with low variability. To evaluate the evolving FDC process, both 4-pt. bend strengths and toughness values are measured on fabricated bars.

Also, the strengths must be reliable and predictable, with low variability. To evaluate the evolving FDC process, both 4-pt. bend strengths and toughness values are measured on fabricated bars. As the FDC process variability has been reduced, strengths have been consistently high (~900 MPa) with large standard deviation (~135 MPa). Fractography was used to identify large (~l 00 /lm) pores as the root cause of low strengths. A Weibull distribution of the strengths, with samples segregated by failure mechanism, illustrates190500Shrinkage during BBOGreen-to-Sintered shrinkage6.0025.004.00:.2.00~ 0.00 ~ -2.00 ·C -4.00 ~ -6.00 '$..

-8.0010.0012.007/4/96II II...• •II? • I • .•...IIII--.II 11~~l1!Iwidth • thickness A length.8/23/96 10/12196 1211/% 1120/97 3/ll197 4/30/9720.00r~ ~ 15.00 ·C ~ 10.00 '$..5.000.007/4/96III 11116 • 1!I •a II II I I •.8/23/% 10/12196 121l/% 1120/97ilL•• 1&•••width II thickness A length3/ll/974/30/97FOCdateFOC dateFigure 6. Application of process improvements decreased linear shrinkage variation to ± 0.5%.the effect of the process improvements (Figure 7). "Vintage 1" samples were made before the topics in this paper were addressed, and "Vintage 2" samples were made as the results of quality control and DoEs were incorporated. "Vintage 2" samples which failed due to large pores have a Weibull modulus (slope of curve) of 4.2, comparable to the 5.6 modulus of "Vintage I" samples.

Application of process improvements decreased linear shrinkage variation to ± 0.5%.the effect of the process improvements (Figure 7). "Vintage 1" samples were made before the topics in this paper were addressed, and "Vintage 2" samples were made as the results of quality control and DoEs were incorporated. "Vintage 2" samples which failed due to large pores have a Weibull modulus (slope of curve) of 4.2, comparable to the 5.6 modulus of "Vintage 1" samples. The distribution of "Vintage 2" samples which broke with no discernible root cause has a Weibull modulus of 15.1, which estimates the process potential if large pores could be eliminated (full sintered density).99.999 99.900 98.000 90.000 70.000 50.000 30.000 20.000 10.000 5.000'C' a:: '-" ~ :a I:L.

Remaining Technical Barriers to Commercializationto commercialization. Large pores must be eliminated from FDC parts to ensure high Weibull modulus and predictable strengths. Internal voids which occur due to insufficient bonding between adjacent roads will be addressed via dynamic, "on the fly" flow control, rather than by less precise off-setting. Sub-perimeter voids remain a problem at the high-incidence angle intersection of rasters and perimeter, as well as at start-stop zones of the raster fills, and flow control will be investigated in conjunction with off-setting.

Statistical Process Control will be instituted to monitor deviation from process targets, and quality control feedback will track each feedstock batch through processing to final sintered part dimensions and strength and isolate aberrant samples.Effortsin the nearfuture willfocus on eliminating remaining obstaclesAcknowledgmentsWe gratefully acknowledge the support of this research by DARPA and ONR under contract # NOOOI4-94-0115, as well as the support and advice of Drs. John Pollinger and Charles Gasdaska of AlliedSignal, Inc., William R. Priedeman of Stratasys, Inc., and Drs. Amit Bandyopadhyay and Gang Qi of Rutgers University.References1.Agarwala, M. K., et. aI., "Structural Quality of Parts Processed by Fused Deposition", Rapid Prototyping Journal, Vol. 2, No.4, 1996, pp. 4-19.2.Agarwala, M. K., et.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperFused Filament Fabrication of Polymer Blends with in situ Layerwise Chemical ModificationsDonald Benza*, Camden Chatham*, Jonathon Baker†, Devin Johnson¶Advanced Engineering Division Savannah River National Laboratory, Aiken, SC†Advanced Technology and Analysis Division Savannah River National Laboratory, Aiken, SC¶Materials Technology and Energy Division Savannah River National Laboratory, Aiken, SCAbstractThe layerwise paradigm of additive manufacturing advertises voxel level control over both geometry and material properties of parts although the latter is difficult to achieve.

Recently the Savannah River National Laboratory has demonstrated a new technique for voxel level material property control via layerwise surface chemical modification of polycarbonate homopolymer with UV and ozone during manufacturing. This technique can be utilized to modify each respective phase of a blended polymer feedstock to increase chemical similarity in preparation for potential interphase crosslinking. Successful crosslinking of dissimilar polymers during in situ manufacturing could allow for further voxel level material property control than modification of a homopolymer could allow. Test feedstock comprised of melt mixed polycarbonate and polystyrene homopolymers, an immiscible polymer blend, were printed in an atmosphere containing ozone and UV light. FTIR measurements indicate both phases of the blend may be modified simultaneously in situ to form new oxygen functional groups, increasing chemical similarity.

FTIR measurements indicate both phases of the blend may be modified simultaneously in situ to form new oxygen functional groups, increasing chemical similarity. increased Calorimetric and compatibilization due to the treatment. Uniaxial tension to failure experiments demonstrates minimal loss of mechanical properties as the blend phase to phase interfacial properties dominates the behavior despite the chemical modifications. Future work will focus on understanding the complex relationships between configurable processing parameters (layer height, print speed, temperature, etc), reaction site creation density, and blend degradation prior to further modification while identifying a suitable crosslinker to improve mechanical and thermal properties of the blend.thermomechanical characterization show noindicators ofIntroductionAdditive manufacturing (AM) advertises voxel level control over both geometry and material properties although, the latter is difficult to achieve.

Recently, the Savannah River National Laboratory (SRNL) has demonstrated a new, uniquely AM approach to voxel level material property control via in situ layerwise surface chemical modifications during manufacture by a reactive atmosphere which is detailed in the authors’ previous work. Surfaces of deposited bisphenol(A) polycarbonate (BPA-PC) layers during Fused Filament Fabrication (FFF) AM were99continuously modified by UV/ozone treatment. Mechanical property changes were observed as a result of the treatment during uniaxial tension to failure testing as UV/ozone modified BPA-PC dogbone samples’ elastic modulus increased ~26% on average while the strain at failure decreased by ~21% on average relative to pristine, unmodified BPA-PC dogbones.

A series of density functional theory (DFT) simulations of reaction between UV/ozone and BPA-PC in the melt state predicted aromatic epoxides and methoxyls would be added to the BPA-PC backbone as the dominant reactions with chain cleavage predicted to be a less probable but additional reaction. Predicted chemical modifications were confirmed by UV/vis and FTIR characterization of the printed samples, validating the addition of oxygen functional groups to the BPA-PC backbone during manufacturing[1]. This technique could be extended to immiscible blended polymer feedstocks such that each phase of the blend could be modified in situ to increase chemical similarity and allow further material property control than in situ modified homopolymer feedstock alone could allow.Blended polymer feedstocks have been successfully processed by multiple AM modalities such as FFF or laser powder bed fusion (L-PBF) to produce parts[2-5].

Traditionally, blended polymer feedstocks have enabled parts to have tunable mechanical, thermal and chemical properties that are difficult or impossible to achieve with commercially available homopolymer feedstock and could bring the same advantages to AM parts. However, a multi-step process is generally required to prepare blended feedstocks prior to manufacturing as many blends are thermodynamically immiscible and exhibit high interfacial tension which results in poor mechanical performance.

Often, compatibilizers, which include block or graft co-polymers, or other chemical modifiers must be manufactured and then added during melt processing of the blend to decrease the polymer A-polymer B interfacial tension and improve polymer dispersion to achieve the desired performance [6].In situ crosslinking of dissimilar polymers during additive manufacturing could increase immiscible polymer blend compatibility and potentially improve material property control at the voxel level allowing for further design flexibility. As a strategy, immiscible polymer blends may be additively manufactured inside an atmosphere containing reactive process gas such as ozone or ammonia to populate each respective phase of the blend with oxygen functional groups such as carbonyls, hydroxyls, and epoxies which could be utilized as common reactive sites for further modification.

Conceivably, a second chemical modifier could then be introduced either as a liquid phase reactant or as a second gas phase reactant to crosslink the new functional groups added to each phase of the blend. Perryman and Dadmun employed a similar strategy to minimize the mechanical anisotropy of additively manufactured homopolymer feedstock. An FFF machine was adapted to spray multi-amines on the surfaces of deposited ABS which readily oxidizes in air above 100 °C. The multi-amines react with the oxidized functional groups to form new covalent bonds which crosslinks the polymer chains located at process induced interfaces resulting in improved fracture energy in the Z-direction [7].

If successful, the strategy outlined above for in situ polymer blend crosslinking could be a cost effective, one step method to create compatibilized polymer blend parts that is applicable to many feedstock formulas as ozone and ammonia are known to react with many types of polymers [8].100In this manuscript, the authors describe the first test case of a blended polymer feedstock additively manufactured in a reactive atmosphere which contains ozone and UV light. The test case feedstock is a filament comprised of a 50/50 weight percent ratio of melt mixed bisphenol(A)- polycarbonate and polystyrene (BPA-PC/PS), an immiscible polymeric blend [9]. The modification of both BPA-PC and PS homopolymer surfaces in the solid state by ozone and UV light treatment has been extensively studied. Each respective polymer forms new oxygen functional groups at similar treatment conditions and timescales. The new oxygen functional groups created are attributed to include carbonyls, hydroxyls, etc.

The new oxygen functional groups created are attributed to include carbonyls, hydroxyls, etc. [10-12].For the remainder of this manuscript, we will first discuss our sample production and characterization methodologies and then the results. An ex vivo polymer-process gas reaction study was setup in an FTIR spectrometer was used to understand both the chemical modifications of the (BPA-PC/PS) blend and respective homopolymers and timescales of the reactions at manufacturing conditions. Dynamic mechanical analysis (DMA) and differential scanning calorimetry (DSC) were used to study thermomechanical and calorimetric changes between the modified and unmodified blended samples. Finally, uniaxial tension to failure experiments were performed to characterize the mechanical performance of both the modified and unmodified blended samples.MethodsFilament Preparation3DX max polycarbonate and polystyrene pellets were purchased from 3DXTech and Sigma Aldrich, respectively.

Pellets were first dried at 80 °C in a vacuum oven prior to processing. The dried pellets were mixed as a 50/50 weight fraction and extruded by a Leistritz ZSE 18. The cooled melt mixed extrusion was pelletized for re-extrusion by a Filabot EX2 configured with a 3.0 mm 3X extended melt nozzle. Melt mixed pellets were re-extruded into 2.85 mm filament at 245 °C and 85% extrusion speed.Printer ModificationsPrinting of all samples was performed by a Ultimaker 2+ Connect configured with an enclosure kit purchased from Matter Hackers. Two VMUS-4 ozone generators supplied ozone at the point of deposition and was measured to be a concentration of 0.9 wt% with a UV-106H ozone analyzer at a flow rate of 6-6.4 SLPM total. Dry air was supplied to the ozone generators via in- line desiccant. Ozone flow was initiated at the start of each print.

Dry air was supplied to the ozone generators via in- line desiccant. Ozone flow was initiated at the start of each print. Two 100 W Everbeam UV lights with a center wavelength of 370 nm illuminated the print bed with a power density of 5 mW/cm2.101Printed Sample PreparationAll samples were printed with the following parameters in Table 1:7 mm s-1 Print Speed 90 °C Bed Temperature 260 °C Nozzle Temperature Layer Height 200 µm Ozone Concentration (Modified Only) 0.9 wt% Ozone flow (Modified Only)6.0-6.4 slpmTable 1: Printed Sample Manufacturing ParametersFourier-transform infrared (FTIR) spectroscopyFTIR intensity measurements of the reaction between filament samples of BPA-PC, PS, and BPA-PC/PS and ozone in time were performed in a Harrick Praying Mantis Diffuse Reflectance Infrared Spectroscopy (DRIFTS) FTIR accessory outfit with the high temperature reaction chamber and setup in a Jasco 6300 series FTIR spectrometer.

A 6.35 x 6 mm cylinder of steel was placed inside of the reaction cup and covered with aluminum foil for filament placement, heat transfer, and ease of sample cleanup. The high temperature reaction chamber dome was outfit with 100 µm thick CVD diamond windows purchased from Applied Diamond in the IR transmission slots. A time series of measurements were made every 40 s. Configurable parameters of the spectrometer were set according to Table 2. All other parameters were set as default/auto.

Configurable parameters of the spectrometer were set according to Table 2. All other parameters were set as default/auto. Filament samples were manually roughed by 80 grit sandpaper prior to placement in the high temperature reaction chamber to encourage diffuse reflectance as opposed to specular reflectance.Scan Speed Scans Interval Time Wavenumber Resolution Aperture0.25 mm s-1 3 40 s 4 cm-1 7.1 mmTable 2 Jasco 6300 FTIR Configurable Parameters for DRIFTS FTIRATR-IR FTIR measurements of printed and filament samples were performed with all settings set to default/auto apart from setting the number of scans to 16 in the same Jasco 6300 series FTIR spectrometer.Differential scanning calorimetry (DSC)Calorimetric characterization by differential scanning calorimetry of sections cut from printed modified and unmodified BPA-PC/PS cylindrical samples of target dimensions 10 mm x 3.2 mm were performed on a Mettler Toledo DSC 3+ differential scanning calorimeter (DSC).

Samples were heated at a ramp rate of 10 °C /min from 0 to 250 °C “as printed”. The glass transition temperature for all samples was calculated by using the midpoint method in the Mettler Toledo STARe evaluation software.102Dynamic mechanical analysis (DMA)Thermomechanical characterization by dynamic mechanical analysis of printed modified and unmodified BPA-PC/PS blend samples was performed on a Mettler Toledo DMA/SDTA 1+ in shear mode. Cylindrical samples of target dimensions 10 mm x 3.2 mm were clamped with 6 in.-lbf of torque and were sheared in oscillation at 1, 10, 100, and 200 Hz. Samples were heated from 70 to 170 °C at a ramp rate of 1 °C min-1. The Mettler Toledo DMA/SDTA 1+ operates in hybrid force/displacement control where one oscillatory cycle is completed either by reaching 20 N of force or 0.1 µm of displacement.Tensile testingASTM D638-14 Type V specification dogbone specimens with grips modified to 12 mm in length were prepared for tensile testing[13].

The specimens were printed in the XY orientation. Samples were printed both in air and in an atmosphere containing ozone and UV light. A MTS Deben Microtest set to a crosshead speed of 0.5 mm s-1 was used to perform uniaxial tension to failure testing.Results and DiscussionSpectroscopic confirmation of in situ chemical modificationBPA-PC, PS, and BPA- PC/PS filaments were monitored for chemical modification by DRIFTS IR over time while at expected processing temperature (90 °C) both with and without ozone exposure, Figure 1,2. All three filament types spectral exhibit no measurable changes after 10 minutes without ozone exposure as expected. Ozone treatment causes changes for all 3 sample types in oxygen regions of the IR spectrum. PS, which contains no oxygen atoms inherently in its backbone, exhibits new, strong peak growth at 1745 cm-1 indicating the addition of C=O bonds which is suggestive of aldehyde formation.

PS, which contains no oxygen atoms inherently in its backbone, exhibits new, strong peak growth at 1745 cm-1 indicating the addition of C=O bonds which is suggestive of aldehyde formation. A small peak growth at 1670 cm-1, a change in C=C stretching, and flattening of peaks at 3058 cm-1 and 3080 cm-1 in the C-H regions are also observed. The growth of these peaks is measurable after 40 s of exposure and likely begins as soon as ozone exposure begins, but changes subside after 80 s of exposure. The formation of C=O bonds on the PS backbone increases chemical similarity with BPA-PC which inherently possesses C=O bonds owed to carbonate groups on the backbone.

The formation of C=O bonds on the PS backbone increases chemical similarity with BPA-PC which inherently possesses C=O bonds owed to carbonate groups on the backbone. BPA-PC exhibits minimal butFigure 1 Photographs of A) Praying Mantis DRIFTS IR setup B) Praying Mantis high temperature Reaction Chamber and C) Unmodified(left) and ozone modified(right) filaments103measurable spectral changes after ozone exposure as a small peak growth at 1670 cm-1 as well as a decrease in peak intensity at 3050 cm-1, is observed. Changes in BPA-PC spectra subside after 80 s of ozone exposure as well. Improved temporal resolution is needed to precisely understand the kinetics of the oxygen functional group formation on each respective polymer.

Improved temporal resolution is needed to precisely understand the kinetics of the oxygen functional group formation on each respective polymer. However, the timescale of reaction propagation observed is well within typical additive manufacturing layer durations.40) .U A.30BPA-PC/PS(e c n a t t i2010BPA-PCm s n a r T01020O3 Modified UnmodifiedPS100015002000 Wavenumber (cm-1)2500Figure 2: DRIFTS IR spectra of PS, BPA-PC and BPA-PC/PS filament both unmodified and ozone modified for 10 minutes in a Harrick Praying Mantis DRIFTS accessory at 90 °CDRIFTS IR spectra of the unmodified melt mixed BPA-PC/PS blended samples exhibit characteristic peaks uncommon to both homopolymers at wavenumbers 1171, 1205, 1257, 1452, and 1787 cm-1.

These peaks are attributed respectively to C-O stretching of carbonate on BPA-PC (1171, 1205 and 1257 cm-1), aromatic C=C stretching in PS (1452 cm-1) and C=O stretching from carbonate ester in BPA-PC (1787 cm-1) indicating chains of both polymer phases are present at the surface and available for modification [12, 14]. Strong asymmetric broadening at the lower wavenumber side of the 1787 cm-1 peak is observed in BPA-PC/PS blended samples but not observed in BPA-PC homopolymer samples. This broadening is attributed to the new C=O bond formation observed in PS homopolymer at 1745 cm-1. Peak growth at 1670 cm-1 and peak flattening at 3058 cm-1 and 3080 cm-1, which was observed for both BPA-PC and PS homopolymer samples alike was also observed in BPA-PC/PS blended samples.

Peak growth at 1670 cm-1 and peak flattening at 3058 cm-1 and 3080 cm-1, which was observed for both BPA-PC and PS homopolymer samples alike was also observed in BPA-PC/PS blended samples. It is likely modified chains from both phases of the blend are contributing to the changes at 1670, 3058, and 3080 cm-1.1043000100) .U A.(e c n a t t im s n a r T80604020020BPA/PC-PS (Printed)BPA/PC-PSPSBPA/PCUnmodified O3 Modified100015002000 Wavenumber (cm-1)25003000Figure 3 ATR-IR spectra of unmodified and ozone modified BPA-PC, PS, and BPA-PC/PS filaments and unmodified and ozone modified printed BPA-PC/PS dogbone specimensBoth filament samples modified in the Praying Mantis high temperature reaction chamber and printed samples’ IR spectra were characterized in ATR-IR mode, Figure 3.

Printed samples could not be characterized by DRIFTS IR due to poor data quality resulting from the relatively smooth surface finish increasing specular reflectance and dominating the signal. Ozone modified and unmodified PS filament samples’ ATR-IR spectra are similar to DRIFTS IR spectra, but BPA- PC filament exhibits new changes in ATR-IR after ozone modification that were not observed during DRIFTS IR, notably broadening of the 1770 cm-1 peak towards lower wavenumbers which was noted by Baker, et al. to indicate aromatic epoxidation. This peak is shifted slightly from 1787 cm-1 in the DRIFTS IR experiments due to the lack of heat. A similar shift is noted for the 1745 cm-1 peak observed in ozone modified PS as this peak shifts to 1735 cm-1 without heat.

A similar shift is noted for the 1745 cm-1 peak observed in ozone modified PS as this peak shifts to 1735 cm-1 without heat. The dissimilar ozone modified BPA-PC spectra between DRIFTS IR and ATR-IR combined with similar spectra for ozone modified PS spectra between the techniques points to an increased diffusion coefficient of ozone in PS relative to BPA-PC as ATR-IR imaging depth is limited to the surface while DRFITS IR can “see” deeper into the bulk[15]. ATR-IR spectra of ozone modified BPA-PC/PS blended filament shows asymmetric broadening towards lower wavenumbers of the 1770 cm-1 peak.

ATR-IR spectra of ozone modified BPA-PC/PS blended filament shows asymmetric broadening towards lower wavenumbers of the 1770 cm-1 peak. Given that both PS and BPA-PC are present at the surface and each modified homopolymer displays change in this region, it is likely modified chains from each phase are contributing to this signal.Like filament samples of BPA-PC/PS, unmodified BPA-PC/PS printed samples exhibit peaks uncommon to both polymers at wavenumbers 1171, 1205, 1257, 1452, and 1770 cm-1105indicating both phases of the blend are present at the surface and available for modification during printing.

Printed ozone modified BPA-PC/PS spectra are very similar to modified BPA-PC/PS filament spectra in the regions 1650 cm-1 to 1800 cm-1 as asymmetric broadening at 1770 cm-1 is observed thus confirming both phases of the blend may be chemically modified simultaneously in situ.Chemical modification impact on blend thermal transitionschemical modifications are made to both phases of the BPA-PC/PS blend, calorimetric and thermomechanical characterization of modified and printed unmodified BPA-PC/PS samples were performed to determine changes in the blend behavior, Figure 4, Calorimetric 3. unmodified characterization printed samples shows two baseline shifts, one at 103.2 °C and a second at 143.4 °C corresponding to the glass transition temperature (Tg) of and BPA-PC phases the PS However, respectively. thermomechanical characterization of the unmodified printed samples displays a single peak at 117.34 °C in the loss tangent corresponding to the polystyrene phase Tg.

thermomechanical characterization of the unmodified printed samples displays a single peak at 117.34 °C in the loss tangent corresponding to the polystyrene phase Tg. Mechanical rigidity of the sample is lost after heating past the Tg of the polystyrene phase and the ordinary polycarbonate and thermal UV/ozone modification of the blend decreases the Tg of the polycarbonate phase by 5.0 °C and 6.3 °C respectively while the polystyrene phase remains unchanged in calorimetric characterization. This suggests one of three scenarios may be occurring. This could indicate dissimilar chain cleave kinetics/energetics between the two phases as chain cleave can explain a decrease in the Tg and chain cleave is a well-known result of ozone modification [8].

This could also indicate that while both phases of the blend are present at the surface and available for modification, the volumetric percentage each phase of the blend is contributing to the surface is dissimilar and thus more BPA-PC chains are modified than PS chains. Lastly, this could alsoGiventhatTableoftherefore masksphase’s behavior. OzoneFigure 4 A) Calorimetric (10 °C min-1, exo up) and B) Thermomechanical analysis of unmodified, O3 modified, and UV/O3 modified printed samples106indicate PS thermal transitions are unaffected by ozone modification. Minimal changes are observed in thermomechanical analysis between modified and unmodified samples regardless of the treatment condition, however, the onset of modulus loss increases from 101.8 °C in the unmodified blend to 104.4 °C in UV/ozone modified samples.

Considering both calorimetric and thermomechanical characterization of modified BPA-PC/PS, there are no strong indicators of increased compatibilization occurring due to the treatment. If this was the case, new peaks which are an average of the thermal properties of both phases or shifting of thermal transitions of each phase closer together would be expected.Tg,a (°C)Tg,b (°C)AirO3UV/O3 AirO3 138.4 N/AUV/O3102.8 DSC DMA 117.6 TABLE 3 Summary of calorimetric and thermomechanical testing results103.2 117.3103.7 117.2143.4137.1Determination of mechanical property changes due to ozone degradationOzone is well known to rapidly accelerate polymer degradation by chain cleavage and other oxidative processes which leads to a significant loss of mechanical properties [8].

In simulations of molten state BPA-PC and UV/ozone reactions performed by Baker, et al., ozone induced chain cleavage of BPA-PC is expected to occur at C-O bonds along the backbone under HOMO-LUMO excitation though it is less probable than methoxylation and aromatic epoxidation. This reaction can account for the decrease in strain at failure observed in UV/ozone modified BPA-PC printed samples and a similar mechanism of action is likely to occur in PS as well. Therefore, a careful balance must be struck when creating potential reaction sites as polymer degradation may also occur.

Therefore, a careful balance must be struck when creating potential reaction sites as polymer degradation may also occur. There may be a tipping point in which too much polymer degradation prior to phase-to- phase crosslinking through further modification could mitigate the benefits to be gained.To begin understanding the effects of UV/ozone pretreatment on the BPA-PC/PS blend mechanical properties, modified and unmodified BPA-PC/PS ASTM D638-14 Type V tensile specimens were printed in the XY orientation (i.e., tensile load applied parallel to deposited roads). Specimens were then elongated to failure in uniaxial tension. Both modified and unmodified specimens exhibit brittle failure modes with minimal material yield prior to failure as is typical in uncompatibilized BPA-PC/PS blends, Figure 5, Table 4 [9]. In unmodified printed dogbone specimens of BPA-PC/PS ultimate tensile strength (UTS) is decreased relative to BPA-PC homopolymer specimens reported by Baker, et al.

In unmodified printed dogbone specimens of BPA-PC/PS ultimate tensile strength (UTS) is decreased relative to BPA-PC homopolymer specimens reported by Baker, et al. with an UTS of 40.3 MPa +/- 0.62 which is expected behavior of uncompatibilized polymer blends. UV/ozone and ozone modified specimens exhibit slightly increased UTS of 42.5 MPa +/- 0.99 and 42.5 MPa +/- 1.48 respectively. The strain at failure is slightly decreased from 0.0198 +/- .001 in unmodified specimens to .0194 +/- .001 in ozone only modified specimens while UV/ozone modified specimens exhibit slightly increased strain at failure, 0.0206 +/- 0.002. The minor changes in mechanical performance between modified and unmodified dogbones suggests that the phase to phase interfacial properties of the blend continue to dominate the performance despite the chemical modifications.

This is not surprising when considering the calorimetric and thermomechanical characterization indicated no107changes in blend morphology or compatibility. Significant changes in the blend mechanical performance will likely not be observed until phase to phase crosslinks are intentionally created by a second chemical modifier which would be expected to improve stiffness, toughness and ultimate tensile strength.Figure 5 Tensile testing of unmodified, ozone modified, and UV/ozone modified BPA-PC/PS printed dogbone blends.

Photographs of B) ozone modified (left) and unmodified (right) tensile bars and C) tensile bars loaded into the MTS Deben Microtest108Specimen Type ModulusUltimate Tensile (GPa) Strength (MPa) 1.795 +/- 0.017 40.29 +/- 0.62 1.965+/ -0.070 42.48 +/- 1.48 1.831 +/- 0.033 42.45 +/- 0.99Unmodified Ozone UV/O3Strain at Failure(mm/mm) 0.0198 +/- 0.001 0.0194 +/- 0.001 0.0206 +/- 0.002TABLE 4: Summary of Tensile Testing ResultsSummary and Future WorkAs this work demonstrates, oxygen functional groups may be simultaneously populated onto each respective phase of BPA-PC/PS blends by ozone surface treatments during additive manufacturing. Ex vivo reaction studies of the interaction of ozone with BPA-PC and PS homopolymers as well as the BPA-PC/PS blend show changes in oxygen regions of IR spectra which conclude after 80 s of treatment. PS forms a new, strong peak at 1735 cm-1 indicating new C=O bond formation increasing chemical similarity to BPA-PC.

PS forms a new, strong peak at 1735 cm-1 indicating new C=O bond formation increasing chemical similarity to BPA-PC. Calorimetric characterization of samples by DSC and thermomechanical testing by dynamic mechanical analysis show no strong indicators of increased interphase mixing or interphase crosslinking occur as a result of the modification. Mechanical testing to failure shows that the treatment does not have a significant impact on mechanical properties as ultimate tensile strength and strain at failure are largely unchanged between unmodified, ozone and UV/ozone treated dogbone specimens as the phase to phase interfacial properties of the polymer blend continue to dominate the mechanical performance. In considering the thermomechanical data collected by DMA, tensile properties uniaxial tensile testing to failure and calorimetric analysis by DSC together, there are no indicators of increased blend compatibilization.

However, the preparation of additively manufactured polymer blend surfaces by ozone to create oxygen functional groups on each respective phase of the blend could allow for further modification that leads to dissimilar polymer crosslinking and increased compatibilization. The newly formed functional groups represent potential common reaction sites populated on each respective phase of the blend that could be crosslinked by a second gas phase reactant or a liquid phase reactant. It will be critical to further understand the effects of the gaseous pretreatment on the blend as there will be complex relationships between configurable processing parameters (layer height, print speed, temperature, etc.), reaction site creation density, and potential blend degradation.A second gas phase reactant or a liquid phase reactant that could crosslink the new oxygen functional groups on each polymer phase will be identified such that properties of the blend may be improved.

Further DFT simulations of PS and ozone will be performed to understand the most probable reaction pathways and precisely identify the functional groups added to the backbone. As it is likely this second reactant will contain nitrogen functional groups and be in the liquid phase, a hybrid FFF/Inkjet system will be used to further this work such that gaseous and liquid reagents may be used simultaneously to crosslink dissimilar polymers in situ. Additional fundamental studies will focus on characterizing the interaction of liquid reagents with the process gas pretreated surface. Critical phenomena to understand will include surface wettability, length109scale of liquid reagent penetration and rheological characteristics of the liquid reagent for jetting control.AcknowledgementsThe authors thank Justin Anderson who supported blend preparation in the twin-screw extruder that was performed at Michael Bortner’s Polymer and Composite Manufacturing Laboratory at Virginia Tech.

Malik Blackman is also thanked for collecting DSC thermograms.This work was produced by Battelle Savannah River Alliance, LLC under contract number 89303321CEM000080 with the U.S. Department of Energy. Publisher acknowledges the U.S. Government the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan). Direct funding was provided through the SRNL Laboratory Directed Research and Development (LDRD) program.licenseto provide public access underReferences1.2.3.4.5.6.7.8.9.10.11.12.Baker, J.N., et al., Layer by Layer Chemical Modification of Polycarbonate- An Introduction to Additive Manufacturing in a Reactive Atmosphere 2023. Chatham, C.A., et al., Semi-Crystalline Polymer Blends for Material Extrusion Additive Manufacturing Printability: A Case Study with Poly (ethylene terephthalate) and Polypropylene. Macromolecular Materials and Engineering, 2019. 304(5): p. 1800764.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceFused Filament Fabrication on the MoonJ. Zhang1, B. Van Hooreweder1, E. Ferraris11Department of Mechanical Engineering, KU Leuven, Leuven 3000, BelgiumAbstractAdditive manufacturing (AM) techniques possess the capabilities to rapidly produce complex and customised parts, typically in low-volume and with reduced material us- age and preparation tools . These attributes magnify their indispensability when the availability of materials and equipment is limited. It can be such a case for off-Earth manufacturing, e.g. in manned lunar explorations. This article presents a preliminary discussion on the possibility of performing fused filament fabrication (FFF) on the Moon from the perspective of heat transfer in printed parts.

This article presents a preliminary discussion on the possibility of performing fused filament fabrication (FFF) on the Moon from the perspective of heat transfer in printed parts. It makes use of experimen- tal data and simulations to quantify the significance of each heat transfer mechanism taking place during printing. The quantification then enables us to investigate how the lunar environment affects the cooling in the printed parts. Finally, FFF on the Moon is predicted to be feasible. Yet, apparent differences in the process window and types of applications are pointed out as compared with the counterpart printing activities on the Earth.The full paper may be found in a special issue of the TMS publication JOM, March 2022.169.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperG-WING: A NOVEL SOFTWARE TOOL FOR TOOLPATH-CENTRIC DESIGN OF WINGS FOR MATERIAL EXTRUSIONJustin D. Valenti∗‡, Joseph Bartolai†, and Michael A. Yukish∗ ∗Department of Aerospace Engineering †Applied Research Laboratory The Pennsylvania State University, University Park, PAAbstractA novel software tool for the design of small aircraft wings to be fabricated with material extrusion is presented where the key requirement of the tool is to minimize the time from identified need to realized capability. The tool, named G-Wing, uses rapid design algorithms based on lifting line theory to determine the outer-mold line of the wing based on desired aerodynamic behavior.

The tool, named G-Wing, uses rapid design algorithms based on lifting line theory to determine the outer-mold line of the wing based on desired aerodynamic behavior. The resulting wing shape and flight-load distribution are given to a structural design algorithm to determine the internal structure of the wing based on both expected flight loads and manufacturing constraints. Finally, manufacturing instructions in the form of G-Code are created directly from the wing shape and internal structure. This process removes explicit geometric modeling and slicing from the critical design path and directly converts airfoil coordinates to perimeter G-Code points, minimizing the introduction of geometric error. This process has been used to design and fabricate multiple small aircraft wings that have successfully flown.

This process has been used to design and fabricate multiple small aircraft wings that have successfully flown. G-Code for an example wing section is shown to be lighter and require less build time compared to G-Code generated by a standard CAD-slicing toolchain.IntroductionSince 2015, a steady line of research exploring the design of Uncrewed Aerial Vehicles (UAVs) to be fabricated with polymer Material Extrusion (MEX) has existed at Penn State [1–8]. From the beginning, it was obvious that the decision to additively manufacture the aircraft changes the design process [1, 2]. Furthermore, there was a desire to rapidly design and iterate on AM UAVs, which lead to development of the design philosphy “Operational Responsiveness” (OpRes) [1].

Furthermore, there was a desire to rapidly design and iterate on AM UAVs, which lead to development of the design philosphy “Operational Responsiveness” (OpRes) [1]. The OpRes philosophy states that the time from identified need to realized capability is minimized when the design space is constrained to the intersec- tion of available product architectures, design tools, and manufacturing methods, shown in Figure 1. “Design assist tools” can be developed for a particular problem which guide and constrain a decision maker produce design within this intersection.The traditional AM workflow involves creating a geometric CAD model, then giving that model to a slicer which translates the geometry into instruction for the AM machine. Presumably, prior to the CAD stage, there are design tools, studies, rules of thumb, and designer’s intuition that are used to generate the design, who’s geometry is then instantiated in CAD. This complete workflow is shown in Figure 2.

This complete workflow is shown in Figure 2. This process works well for many applications. Aircraft wings are not one of these applications.‡Corresponding Author: jdv5076@psu.edu700Figure 1 : The idea of designing for operational responsiveness focuses on constraining the design space to the intersection of available manufacturing methods, design tools, and product architectures.Figure 2 : Traditional AM WorkflowThe strict strength to weight requirements of aircraft structures, combined with the minimum feature size of an MEX machine, tend to lead to single extrusion sparse structures. Not only are these types of structures difficult for MEX machines to fabricate, slicers struggle to generate a “good” path plan for the single extrusion structures. The wing geometry needs to be generated with an understand of how the final G-Code is generated to ensure that the final wing matches the designer’s intent. Two publicly available examples demonstrate this.

Two publicly available examples demonstrate this. First, the STL files of the freely available Eclipson Airplane [9] have strategically place701voids within the wing structure. This manipulates the slicer into creating a part only with single extrusion walls. The wing is designed to be sliced with 0% infill and for the perimeter extrusions to form both the skin of the wing and the internal structure. Second, the company 3D Lab Prints [10] sells the files to build aircraft on Desktop 3D printers. They provide the geometry in several file types and specific instructions on how to slice the geometry in several available slicers. Both examples show clear engineering thought into not only the shape of the aircraft, but how the aircraft is converted into G-Code.In aircraft structures, failure to invest in the toolpath design will likely lead to a poor end product.

It is important to consider upfront when designing a wing and its internal structural how that structure will be translated into a coherent toolpath. In short, for MEX-produced wings, the wing design problem is really a toolpath design problem. The rest of this paper details a fast, efficient, and holistic solution to that design problem.Novel Contribution: A holistic toolchain was developed to design UAV wings to be built with MEX. This toolchain was instantiated in a software tool named G-Wing. The software tool connects low-order aerodynamic, structural, and process design algorithms to automate the design process from desired aerodynamic behavior to ready-to-make G-Code that guar- antees flight ready wings.

This process bypasses the CAD and slicing stages of the tradition AM toolchain, minimizing the introduction of errors, and ensuring that the designers intent is carry all the way to the finished product.Layout of Paper : This manuscript first provides a detailed description of the G-Wing functionality and underlying design modules. We then provides a comparison of the G-Wing toolchain to a more conventional MEX toolchain. A discussion section explores the broader implications of this work. Finally, a conclusion section summarizes the work.Description of G-WingG-Wing is a custom software tool written in Python 3. G-Wing designs wings using a recently publish design pattern specifically developed for wings build with MEX [11]. The tool consisted of 3 main modules: aerodynamics, structures, and process planning. The tool ingests a text-based configuration (config) file in which the user defines all aerodynamic, structural, and process design parameters and settings (Figure 3).

The tool ingests a text-based configuration (config) file in which the user defines all aerodynamic, structural, and process design parameters and settings (Figure 3). The main output of G-Wing is a collection of .gcode files ready to be given to an MEX machine.Aerodynamic Design: The main output of aerodynamic design module is the outer mold line of the wing and the spanwise lift distribution across the wing. Lifting line theory [12,13] provides a fast method to analyze wings. The authors have publish several algorithms [14,15] to perform the inverse design of wings using the matrix formulation [16] of lifting line theory. These algorithms are used in G-Wing to allow the user to generate wings to yield a desired aerodynamic behavior. A brief overview of the model will now be provided.In lifting line theory a 3D wing is modeled as 1D straight “lifting line” which runs in the spanwise (y) direction from the left wingtip to the right wingtip, as shown in Figure 4.

This formulation is valid for straight wings with no sweep or dihedral. The outer mold line of the wing is uniquely defined by an airfoil shape∗, a continuous distribution of chord as a function of y, ˜c = ˜c (y), and a continuous distribution of twist of a function of y, ˜α = ˜α (y). In the∗The present work assumes a constant airfoil shape, but the work is easily generalized to allow for avarying airfoil across the span.702Figure 3 : Workflow of G-Wing design modulesmatrix formulation, the lifting line is discretized into n equally spaced spanwise stations. The chord and twist distributions are then approximated a n × 1 column vectors where each element of the vector represents of local value of chord or twist: ˜c (y) ≈ ⃗c = [c1, c2, . . . , cn]T , ˜α (y) ≈ ⃗α = [α1, α2, . . . , αn]T . This discretization is shown in Figure 4.the cross section of the wing) is represented by a list of (x, z) coordinates.

. . , cn]T , ˜α (y) ≈ ⃗α = [α1, α2, . . . , αn]T . This discretization is shown in Figure 4.the cross section of the wing) is represented by a list of (x, z) coordinates. This list is stored as a m × 2 matrix, where m is the number of airfoil coordinates: Paf = [x1, z1; x2, z2; . . . ; xm, zm]. G-Wing does not have the ability to perform airfoil design. The airfoil shape is provided up from by the user.While not pictured, the smooth airfoil shape (i.e.In the matrix lifting line formulation, the wing is represented by n adjacent “horse shoe vortices”. Each horseshoe vortex is fixed to the lifting line, centered about a y location, with tails trailing off infinitely far downstream, as shown in Figure 4. At an arbitrary y location, the local lift ℓ is directly related to the vorticity Γ at that location via the Kutta-Jouskowski theorem, ℓ = ρV∞Γ.

At an arbitrary y location, the local lift ℓ is directly related to the vorticity Γ at that location via the Kutta-Jouskowski theorem, ℓ = ρV∞Γ. Following the discretization scheme of chord and twist, the continuous lift and vorticity distributions are also approximated as n × 1 vectors: ⃗ℓ = [ℓ1, ℓ2, . . . , ℓn]T and ⃗Γ = [Γ1, Γ2, . . . , Γn]T , respectfully.Several design algorithms are available to the user in G-Wing. Of which, the user chooses one to design the wing. This first design algorithm use the basic lifting line calculation [16] which calculates a vorticity distribution ⃗Γ from the wing geometry: ⃗c and ⃗α. This allows a designer to explicitly define the outer mold line up front. G-Wing then calculates the aerodynamic behavior.The next two algorithms, formally presented in [14], allow a user to specify a lift dis- tribution ⃗ℓ and either a chord distribution ⃗c and a twist distribution ⃗α.

G-Wing will the calculate either ⃗c or ⃗α, whichever was not specified.703Figure 4 : Diagrams of the definition of external wing geometry (left) and how the wing is modeled in lifting line theory (right). Lifting line theory collapsed the 3D wing geometry into a 1D lifting line.It is worth noting that the lift distribution will change as the wing changes total lift coefficient, defined as cL = 2L/ (ρ∞V 2 ∞S), where L is the total lift over the entire wing, ρ∞ is the freestream air density, and S is the wing area. Wings operate over a range of cL values, typically trading off between V∞ and cL. High cL values correspond to slower flight and vice versa. The prior three designs allow a user to design a wing for the aerodynamics of a single cLFinally there are three algorithms that allow the user to design a wing by defining the aerodynamic behavior at two cL values.

The first, presented in [14], allows a user to define two lift distributions, ⃗ℓ1 and ⃗ℓ2, each at a different cL, and simultaneously calculate the wing chord and twist. The final two algorithms, presented in [15], involve designing a particular spanwise distribution of local 2D lift coefficients ⃗cℓ = [cℓ1, cℓ2, . . . , cℓn]T . At the jth spanwise location, the 2D lift coefficient is defined as cℓj = 2ℓj/ (ρV 2 ∞cj). Algorithms are available to calculate chord and twist from (1) two local lift coefficient distributions, ⃗cℓ1 and ⃗cℓ2, each at a different cL and (2) one lift distribution, ⃗ℓ1, and one lift coefficient distribution, ⃗cℓ2.A summary of all available aerodynamic design algorithms is provided is in Table 1. Formal mathematical derivations and instructions on how to implement the algorithms are704provided in [14, 15].

Formal mathematical derivations and instructions on how to implement the algorithms are704provided in [14, 15]. Regardless of which algorithm is used, at the end of the aerodynamic design routine, the wing geometry is full defined by the a chord distribution ⃗c, a twist distribution ⃗α, and a user defined airfoil. The flight loads are defined by the lift distribution ⃗ℓ. As a note, in the case of designing a wing for two cℓ distributions, the results need to be run through the direct design algorithm to obtain a lift distribution.Table 1: Six design algorithms are available in G-Wing, of which the user would chose one to perform the design.

The first three algorithms design a wing for a single cL, the second three algorithms allow a designer to define the aerodynamic behavior at two different cL values.Design CaseUser Defined Inputs Output(s)Direct Design Single Point Inverse Design #1 Single-Point Inverse Design #2 Two-Point Inverse Design #1 Two-Point Inverse Design #2 Two-Point Inverse Design #3⃗c, ⃗α ⃗c, ⃗ℓ ⃗α, ⃗ℓ ⃗ℓ1, ⃗ℓ2 ⃗cℓ1, ⃗cℓ2 ⃗ℓ1, ⃗cl2⃗ℓ ⃗α ⃗c ⃗c, ⃗α ⃗c, ⃗α ⃗c, ⃗αSince the lifting line model collapses the wing to a straight line, there is some ambiguity on how non-linear chord and twist distributions are converted into 3D geometry. This is handled in G-Wing by aligning all airfoil sections about a potential hinge line for a control surface. The chordwise location of the this straight line, and thus the resulting hinge-line, is a user-defined parameter.

The chordwise location of the this straight line, and thus the resulting hinge-line, is a user-defined parameter. The conversions chord and twist distributions into a wing planform is shown Figure 5.It is interesting to note that the outer mold line of the wing is fully defined by the chord and twist distributions, both in the form of n × 1 vectors, and an airfoil profile in the form of (x, z) coordinate list. If the wing is discretized into 400 spanwise stations and the airfoil coordinate list contains 40 (x, z) points, both reasonable values, then the entire wing geometry is defined by 880 floating point numbers. At this point there is no explicit geometric modeling or tessellation because it is not needed at any point from design to G-Code.G-Wing can optionally call Athena Vortex Lattice (AVL) [17] for aerodynamic verifica- tion. If the designer opts for AVL comparison in the config file, G-Wing will output the wing geometry to AVL for higher-order aerodynamic analysis.

If the designer opts for AVL comparison in the config file, G-Wing will output the wing geometry to AVL for higher-order aerodynamic analysis. The aerodynamic behavior predicted from the lifting line theory of G-Wing is then automatically compared to the behavior predicted by the vortex lattice theory of AVL.Structural Design: Given the outer mold line, the structural design model determines the layout of the internal structure. Previous work by the authors’ [4] propose a family of structure known as “curvilinear spars” that allow the internal structure to be tailored for a desired bending stiffness. The concept was inspired by work on “spaRibs” from Virginia Tech [18] and incorporated the concept of “cross-sectional analysis” [19] to simplified the design process.The wing is built “standing up” with the build direction oriented with the spanwise direction. Each layer consists of the airfoil shape and several vertical members, i.e spars, as shown in Figure 6.

Each layer consists of the airfoil shape and several vertical members, i.e spars, as shown in Figure 6. At any given spanwise location, the number and chordwise (x) locations705Figure 5 : Example wing (bottom) defined by chord and twist distributions (top two plots). The dashed line in the bottom plot corresponds to a straight hinge line at 80%chord.Figure 6 : Sectional view showing an airfoil with several spars at various chordwise (x) locations. The “main” spar is placed at the airfoil max thickness location xτ max. The number and locations of the spars fore and aft of the main spar are determined by bending inertia requirements and AM constraints.of the spars chosen to either increase the bending inertia of the wing to carry the flight706loads [4] or support the skin such that it does not deform during manufacturing [7, 8]. The internal structure of the wing can then be uniquely defined by an m × n “spar location matrix”:xsp11 xsp21 . ..

The internal structure of the wing can then be uniquely defined by an m × n “spar location matrix”:xsp11 xsp21 . . . xspm1 xspm2xsp12 xsp22 . . .. . . . . . . . . . . . xspmnxsp1n xsp2n . . .    Xsp =.A single row of the Xsp matrix represents the chordwise locations of a single spar across the span. A single column represents the chordwise locations of all spars at a single spanwise location. The reader is directed the previous cited work for the exact details on of the these two design considerations, but a brief pseudo-code of the design process is given in Algorithm 1.Algorithm 1: Pseudocode describing the steps to generate the spar locations form a wing load, wing geometry, and AM constraintsInput: Wing Design: {Chord distribution ⃗c, Lift distribution ⃗l,Airfoil Coordinates Paf }Machine Settings: {extrusion width, material yield strength σy}Output: Spar Location Matrix Xsp1 Locate the chordwise location of the max airfoil thickness.

This becomes thelocation of the “main spar”;2 Calculate the bending moments on the wing from the lift distributionprovided by the aerodynamic module;3 Calculate the bending inertia required to carry the bending moments; 4 Place spars to provide the required bending inertia according to [4]; 5 Limit spar spacing based on AM limits based on [8]; 6 Write the spar locations to the matrix XspTwo example wing structures are provided in Figure 7. The structure shown in Fig- ure 7(a), is example wing structure that was generated to only carry the flight loads with no consideration of supporting the wing skin. This corresponds to the structure that is generated by Step 4 of in Algorithm 1. The structure given in Figure 7(b) shows the type of structure that results from taking the structure shown in 7(a) then limiting the max spar spacing to properly support the wing skin. This corresponds to the structure that results from Step 5 in Algorithm 1.

This corresponds to the structure that results from Step 5 in Algorithm 1. It can be seen in the aft portions of the wing, around spanwise locations of roughly ±0.15y/b, that the curvature of the spars distinctly changes. Inboard of this location, spar locations are fully designed to carry flight loads. Outboard of this region, the fore spars are still designed to carry the flight loads, however the aft spars are designed to support the skin during manufacturing.The main output of the structural design module is the spar spacing matrix Xsp.

This matrix, along with the chord and twist distributions found in the aerodynamics module fully707(a) Planform view of curvilinear spar structure designed for a particular bending inertia to carry the expect flight loads(b) Planform view of curvilinear spar structure with the chordwise spar spacing limited by AM constraintsFigure 7 : Planform view of example curvilinear spar structures both (a) without and (b) with AM spar spacing constraints. In both structures, the black dashed line corresponds to the “main” spar placed along the max thickness points of the airfoil sections. The remaining dashed lines correspond to the other curvilinear spars.

The remaining dashed lines correspond to the other curvilinear spars. Also visible is the hinge line running between the wingtips (explained in the Aerodynamics section) and the wing section breaks (explained in the Process Planning section).define the shape of the wing and are provided to the process planning module to generate the G-Code.Process Planning: With the outer mold line defined in the aerodynamic module and the internal structure defined in the structural module, the process planning module is tasked with generating the instructions to the MEX machine that represents the final design of the wing. MEX parameters such as build volume dimensions, extrusion width, layer height, build speeds and temperatures, etc. that are user-defined in the config file are used by the process planning routine to translate the airfoil, chord, twist, and spar spacing distributions into build-able G-Code files.Wingspans are typically too large to be fabricated in one build.

The process planning module automatically breaks the wing into smaller sections based on build volume. The wing sections are built “standing up”, with the spanwise direction of the wing corresponding to the build direction of the MEX machine†, as shown in Figure 8. In this orientation, the span of a single wing section is limited by a user-defined max build height,zmax. This parameter can either correspond to the height of the build volume itself or, more likely, the max build height that can be achieved before the tall, thin wing section begins to wobble. The first†Please note that standard aircraft coordinates define the spanwise direction as the y direction. In AM literature, the build direction is referred to as the z direction.

The user defines a center section width wc, which could correspond to the width of a fuselage. After the center section is defined, the remain wingspan is broken into wing sections with a span ofbsect =(b − wc) ceil ((b − wc)/zmax).An example of these wing section breaks are shown in Figure 7.With this build orientation, the perimeter points for each layer corresponds to the airfoil coordinates. The size and angle of the airfoil is provided by the chord and twist, which was determined in the aerodynamic design module. The infill locations are determined by the spar spacing found in the structural design module. Example tool paths are shown in Figure 8. One exception to the infill is that solid infill (i.e. spars with no space in between) is used to the first layer of each wing section. This both improves the build plate adhesion and serves as a wing rib in the final structure.

spars with no space in between) is used to the first layer of each wing section. This both improves the build plate adhesion and serves as a wing rib in the final structure. Pseudocode for the G-Code generation is provided in Algorithm 2.The process planning module outputs a series of G-Code files, one for each wing section. These G-Code files represent the final design of the wing.

Nonetheless, CAD modeling can still be useful when incorporating the wing into a broader aircraft. G-Wing can optionally generate an OpenSCAD [20] model of the wing. Notably, the process to generate the .scad file directly from the aerodynamic and structural design results and is fully independent of the G-Code generation process.The independence of CAD and G-Code generation means that both are theoretically exact representations of the intended design. One is not a interpretation of the other. There is no loss of designers intent on the path to G-Code.OpenSCAD was chosen as the CAD software for two reasons. First, it is an open-source software that is freely available online. Second, the file format is purely text-based. As a result, G-Wing needs to only generate an additional text-based .scad file.

First, it is an open-source software that is freely available online. Second, the file format is purely text-based. As a result, G-Wing needs to only generate an additional text-based .scad file. The user would then have option to open this .scad file in the OpenSCAD after G-Wing finishes running.710This meaning that that CAD output does not interrupt the design flow.Comparison to Conventional Tool ChainA comparison of G-Code generated with G-Wing to that generated with a more conven- tional AM toolchain was performed to explore the benefits of G-Wing. A wing section was chosen for this comparison from a wing designed for the Kite, shown in Figure 9, a small UAV designed as a test-bed for MEX-built wings. The wing itself has a factor of safety of greater than 20 and the spar spacing is determined by manufacturing constrains everywhere along the span.

The wing itself has a factor of safety of greater than 20 and the spar spacing is determined by manufacturing constrains everywhere along the span. The wing and selected test section is shown in Figure 10.(a) Kite with 3 wings of different designs(b) Kite in flightFigure 9 : The Kite is an MEX-produced UAV, which was designed as a test-bed for MEX wings.(a) Planform view of wing with test section boxed in red(b) G-Code of test section generated by G-WingFigure 10 : Test wing section chosen for study.A summary of the wing parameters is provided in Table 2. The MEX parameters used to generate the G-Code is provided in Table 3. An OpenSCAD model of the test section was also generated by G-Wing, which was in turn exported as an STL file. The open-source slicing software Slic3r [22] was used to slice the wing section STL with several common infill settings.

The open-source slicing software Slic3r [22] was used to slice the wing section STL with several common infill settings. G-Code files were generated with 5% and 20% infill fractions for grid, honeycomb,711Table 3: Build ParametersTable 2: Wing parametersWing Parameter Airfoil Wingspan Aspect Ratio Spar SpacingValue AG24 [21] 1.09 m 8 30 mmFFF Parameter Nozzle Diameter Layer Height Extrusion Width Material Density (PLA) Perimeter Speed Infill Speed Solid Infill Speed 1st Layer Speed Travel SpeedValue 0.4 mm 0.2 mm 0.5 mm 1240 kg/m3 40 mm/s 40 mm/s 60 mm/s 30 mm/s 80 mm/sand aligned rectilinear sparse infill patterns. The build parameters listed in Table 3 were also used for slicing in Slic3r for an apples-to-apples comparision with G-Wing. Screenshots of the 6 G-Codes generated with infills in Slic3r are provided Figure 11.

Screenshots of the 6 G-Codes generated with infills in Slic3r are provided Figure 11. A seventh G-Code(a) Grid 5% Infill(b) Honeycomb 5% Infill(c) Aligned Rectilinear 5% Infill(d) Grid 20% Infill(e) Honeycomb 20% Infill(f) Aligned Rectilinear 20% InfillFigure 11 : Wing sections sliced in Slic3r with infill.file was generated in Slic3r with no infill as a baseline to show the effects of infill on the build.Both G-Wing and Slic3r predicts part mass when G-Code is exported. The build times of wing sections were predicted using the open-source software Pronterface [23]. The predicted mass and build times of all G-Code files are provided in Table 4. Considering mass first, G-Wing is about halfway between the three 5% infill sections and the 0% infill section. This immediately indicates the how sparse these wing structures are. Slic3r’s manual suggests infills of at least 20% for any part where strength is important.

This immediately indicates the how sparse these wing structures are. Slic3r’s manual suggests infills of at least 20% for any part where strength is important. Unfortunately, the 20% infill sections over twice the weight of the G-Wing sections.The build time comparison is also very interesting. The G-Wing section is only marginally longer than the 0% infill section. This highlights the efficiency of the G-Wing toolpath. By712Table 4: Mass and build time values of the wing section generated by G-Wing and wing sections generated by Slic3rMetricMass (g) Build Time (h:mm)G-Wing36.2 2:10Slic3r 0% Infill 24.30 2:07Grid44.38 4:11Slic3r 5% InfillHoney- Comb 48.20 4:18Aligned Rect. 43.69 3:14Slic3r 20% InfillGrid69.76 5:39Honey- Comb 78.67 6:12Aligned Rect. 70.28 4:53contrast, the other wing sections take hours longer to print.

43.69 3:14Slic3r 20% InfillGrid69.76 5:39Honey- Comb 78.67 6:12Aligned Rect. 70.28 4:53contrast, the other wing sections take hours longer to print. Notably, the aligned rectilinear sections require significantly less time than the other wing sections of similar infill percentage. This suggests aligned rectilinear may be a good choice of infill if a designer is required to use a conventional tool chain.Finally, it is worth looking qualitatively at the lattice infills. The G-Wing design is strong enough to carry the flight loads, which suggests infill percentages of any infill design for this section ought to be around 2-3%. At these infill percentages, an individual lattice cell does not fit within the wing section. The sparse lattice infills become a non-uniform collect of inefficient extrusions.DiscussionFrom the AM-specific prospective, G-Wing solves several problems that appear in a typ- ical AM toolchain.

First, there is no loss in designer intent throughout the entire toolchain. In a tradition CAD-slicer toolchain, there are opportunities for the introduction of geometric error first when modeling the design in CAD, then again when the design is tessellated into an STL, then again when that tessellated shape is translated into G-Code. The development of G-Wing began by asking the question “what does a good toolpath look like?” Our answer to that question is provided in Figure 8. From there, the rest of the design methods were developed and link together in a way maintained the designers intent through the entire process.A second AM-specific implication of G-Wing is that the process planning is completely deterministic. The same config file with always lead to the same G-Code. During a tradi- tional slicing session, it is often difficult to know and/or decide exactly how infill patterns will be placed and exact what the pathplan will be.

During a tradi- tional slicing session, it is often difficult to know and/or decide exactly how infill patterns will be placed and exact what the pathplan will be. G-Wing solves that problem.The restriction of this work is that G-Wing is application specific - the inevitable result of the focus on extremely short design cycles as the key performance metric. The software will only produce wings within a very specific design space and nothing more. Any addi- tional feature is only achieved through some amount of development effort. For a different application, the decision has to be made whether or not the specific design problem at hand is repeated enough times to warrant a full engineering effort to develop the digital thread. This leads to the broader digital engineering implications of the work.There is currently a large amount of enthusiasm surrounding “digital engineering” [24, 25].

This leads to the broader digital engineering implications of the work.There is currently a large amount of enthusiasm surrounding “digital engineering” [24, 25]. A major concept in digital engineering is taking ownership of the toolchain of an engineering problem. G-Wing is a small example of a homegrown, fully-owned toolchain from initial design to complete manufacturing process plan. For another application, it713may make more sense to integrate a combination of homegrown and commercial-off-the- shelf tools. Regardless, the building and owning of that toolchain is a digital engineering problem. When this is done properly, the time from identified need to realized capability is minimized. G-Wing is an example of this - the tool itself designs a ready-to-build wing less than 5 minutes that will fly.ConclusionG-Wing is a holistic toolchain which automates the design from desired aerodynamic behavior to ready-to-build G-Code.

The software tool allows a designer to simultaneously perform aerodynamic, structural, and process design. The resulting G-Code was shown to be significantly better than G-Code created by a more traditional toolchain using standard slicing settings.At the time of writing this manuscript, the source code itself for G-Wing is currently undergoing public release. A GitHub repository [26] has been set up in anticipation of public release. The code is expected to be published before this manuscript is available online.References[1] S. W. Miller, M. A. Yukish, M. E. Hoskins, L. A. Bennett, and E. J. Little, “A Retro- spective Analysis of System Engineering Data Collection Metrics for a 3D Printed UAS Design,” in Procedia Computer Science, vol. 153. Elsevier B.V., 2019, pp. 1–8.[2] S. W. Miller, M. E. Hoskins, L. A. Bennett, E. J. Little,and M. A. Yukish, “CHALLENGES AND LESSONS LEARNED FROM DEVELOPING AN OPERATIONALLY RESPONSIVE 3D PRINTED UAV,” Penn State University, Tech.

INING COMPETITIVE ADVANTAGE BY USING RAPID PROTOTYPING IN AIRCRAFT MODIFICATIONAuthors: Mr. Steven L. Smith ManufacturinglProducibility Engineer, Raytheon-WacoMs. Jerri Paul Process and Management Consultant, Raytheon-Waco Masters Candidate, University of Texas at Austin1.0 INTRODUCTIONRaytheon-Waco (formerly Chrysler Technologies Airborne Systems) is an aircraftmodification facility in Waco, TX, specializing in military, executive, and head-of-state aircraft modifications. Raytheon-Waco has over 25 years of aircraft modification experience, and is a leading contractor for major US and foreign government programs that require system design, airframe modification, installation, flight testing, and post-delivery technical and logistical support.

Current programs at Raytheon-Waco include foreign government Head-of-State wide body aircraft modifications consisting of luxury interiors, secure communications, and airframe modifications, and US and foreign military aircraft modifications involving electronics, communications, avionics, and airframe modifications.The aircraft modification industry is highly competitive, with a variety of very competent companies constantly trying to increase market share and move into new markets. Raytheon Waco has focused primarily on military communications and transport aircraft, and Head-of State wide-body aircraft. Due to increasing competition and customer cost-conscientiousness, Raytheon-Waco is working to reduce cost and cycle time while maintaining its reputation for delivering on time and within budget.

Due to increasing competition and customer cost-conscientiousness, Raytheon-Waco is working to reduce cost and cycle time while maintaining its reputation for delivering on time and within budget. There is also movement into other markets, such as lower cost wide-body executive aircraft.Raytheon-Waco is currently transitioning to a paperless design process, with the ultimategoal of performing all design, analysis, manufacturing, and aircraft installation from CAD models without requiring paper drawings. Increased use of electronic models will also allow more parallel processes to be performed, eliminating many of the sequential design steps currently required. Among the steps that will be performed in parallel will be detailed design, manufacturing and installation planning, technical publications, and design analysis. Without improved initial designs, parallel work will only lead to increased rework and wasted effort.

Without improved initial designs, parallel work will only lead to increased rework and wasted effort. Rapid prototyping is a key factor in reducing risks associated with concurrent processes, and its benefits are critical to efforts to reduce the cost and schedule of future aircraft modifications.2.0 RAPID PROTOTYPING AT RAYTHEON·WACO 2.1 RAPID PROTOTYPING NEEDTraditional high-volume production is seldom required at Raytheon-Waco. Most of thecommercial wide-body and military modifications are unique designs that are similar to traditional prototypes in that the designs are constantly evolving, typically only small pieces and57sub-assemblies of previous designs can be re-used, and lessons learned in previous aircraft modifications only have limited application.

Reducing cost and cycle time for many aircraft components by quickly and cheaply producing prototypes of new designs for customer approval and design verification on the aircraft prior to part fabrication was a real need, and was the driving factor behind the acquisition of rapid prototyping (RP) equipment. Potential manufacturing and tooling uses of RP technologies also contributed to the decision to invest in RP, but were not the driving factors.2.2 RAPID PROTOTYPING EQUIPMENTIn 1995, Raytheon-Waco began investigating rapid prototyping equipment to be usedprimarily by design engineers to ensure form, fit, and function of parts and assemblies. This led to the purchase of a 3D Systems SLA-500 Stereolithography Apparatus (SLA) in early 1996. In the SLA process a solid physical part model is made directly from 3-D CAD model data within a relatively short period of time, usually less than one day.

In the SLA process a solid physical part model is made directly from 3-D CAD model data within a relatively short period of time, usually less than one day. The model is built by successively laser curing layers of photopolymer resin until the part is completed. After the model has been built, it is removed from the SLA processing chamber and placed in a Post Curing Apparatus (PCA), where the model is flooded with intense ultraviolet light to complete the polymerization process. It is then finished using a variety of sanding and painting techniques. The total processing time is typically one to three days.2.3 SIA IMPLEMENTATIONRaytheon-Waco has a laboratory dedicated to rapid prototyping, with the SLA-500, an accompanying centrifuge and PCA, as well as a model finishing facility. The SLA technician is responsible for finishing each model, including post-curing, removing excess resin, sanding, painting, and when necessary bonding models into assemblies.

The SLA technician is responsible for finishing each model, including post-curing, removing excess resin, sanding, painting, and when necessary bonding models into assemblies. The SLA technician works directly with engineers considering SLA models to determine if they have a good application, modify CAD models to improve the SLA model quality, and ensure that the part model can be built in the SLA-500. He will also input the appropriate shrink factors into the build process for models used as tooling.SLA modeling is part of a move to paperless design, fabrication, and installation in order toreduce costs and compress schedules. The process of ordering an SLA model is completely electronic, and requires 3-Dimensional (3D) surface or solid models. New design work at Raytheon-Waco is done in ProlEngineerl 3D solid modeling, which has a standard .STL file output that can be input into the 3D Systems software.

New design work at Raytheon-Waco is done in ProlEngineerl 3D solid modeling, which has a standard .STL file output that can be input into the 3D Systems software. There is also an approved ISO 9001 procedure and work instruction for electronically ordering SLA models.3.0 SLA IN DESIGNIn the six months since obtaining RP capability, Raytheon-Waco has realized significantbenefits from using SLA, with potential for even larger benefits in the future. Specific advantages have been seen in the design process and in an increase in customer input and satisfaction. Allowing the designers, engineers, management, purchasing, and customers to review physical models of designs before releasing a drawing is expected to reduce the number of engineering change orders issued during the design and production phase. In addition, check fitting full scale SLA models of various parts and assemblies in the aircraft is expected to reduce58last minute changes resulting from airframe variations.

In addition, check fitting full scale SLA models of various parts and assemblies in the aircraft is expected to reduce58last minute changes resulting from airframe variations. These applications will in turn reduce the amount of lost production time, retooling, rework, and scrap waste incurred during manufacture and installation. Following are some specific examples of uses of SLA in the design process.3.1 INCREASING CUSTOMER INVOLVEMENT IN DESIGN 3.1.1 Aesthetic ReviewOne of the fIrst uses of SLA was in producing models of artistic complex-curved surfacesused in luxury interiors prior to producing part drawings or beginning design work on many interfacing components. These models were reviewed with the customer, who determined that the initial design needed to be changed.

These models were reviewed with the customer, who determined that the initial design needed to be changed. Without rapid prototyping, these design changes would not have been made until much later, and possibly not until the parts had been fabricated and significant design had been done on interfacing parts, resulting in additional design changes. SLA models have become common elements of customer reviews, beginning with the earliest stages of the design process.3.1.2 Functional ReviewsSLA models were made of electronic control panels to demonstrate proposed Air Force C141 transport cockpit upgrades. The panels were approximately 5" x 3" and showed the expected configurations of meters and switches. The SLA panels were assembled into an on-site cockpit simulator, allowing Air Force customers, including pilots, to provide specific input on changes to increase ergonomics and functionality of the new cockpit elements.

Previously, only 2-D drawings were available at this stage in the design process, and cockpit mock-ups were only available much later.SLA models of several critical parts of a system being integrated into a military aircraft weremade for design review meetings with the primary contractor and Air Force representatives. Scale models were reviewed in a series of meetings and several design changes were made. Program officials estimated that without the SLA models, the changes would not have been made until much later at a cost of about $10,000 in time and materials. The same program also produced full-scale SLA models of lofted contoured panels that were taken to the aircraft to confrrm existing data and determine fastener locations.

The same program also produced full-scale SLA models of lofted contoured panels that were taken to the aircraft to confrrm existing data and determine fastener locations. The SLA models used in this program have saved many times their expense in time and materials by avoiding design changes later in the project.3.2 AIDING THE DESIGN PROCESS 3.2.1 Aiding Design EngineeringEngineers at Raytheon-Waco have been encouraged to use SLA models to better visualize complicated designs and confrrm that the design will meet their requirements. Some engineers have had models made of small parts they are designing so that they could take them to the aircraft to confrrm form and fit of the part before proceeding with additional design work.

Some engineers have had models made of small parts they are designing so that they could take them to the aircraft to confrrm form and fit of the part before proceeding with additional design work. These same parts are then available to other engineers considering similar designs, or designing parts that must interface with the modeled parts.An SLA model was made of an exit duct and a machined fitting for a galley being redesignedfor a foreign Head of State aircraft. The fitting was a Y-shape with complex contoured flanges.59The model was used by the design engineers and the customer to identify interface requirements and recommend design changes to increase manufacturability and serviceability. The exit duct SLA model was used for a producibility review, and will become a washable sand mold to make a fiberglass part.

The exit duct SLA model was used for a producibility review, and will become a washable sand mold to make a fiberglass part. The SLA model will increase the manufacturability of the part and allow the manufacturer to decrease the lead-time and costs associated with a prototyPe part.3.2.2 Verifying Designs with SLA ModelsIn order to verify fit with existing plumbing interfaces and avoid interference with otherstructures, SLA models of a custom plumbing system were created for a commercial VIP lavatory refurbishment. The SLA models were installed in the aircraft and revealed interference problems that would not have been detected until much more costly metal parts were fabricated later in the program. New models of design changes were also produced to verify all changes prior to creating drawings. These models will also be available for other engineers to use to ensure that other systems proPerly interface or avoid interference with the lavatory plumbing fixtures.

These models will also be available for other engineers to use to ensure that other systems proPerly interface or avoid interference with the lavatory plumbing fixtures. It is eXPected that the total savings from using plumbing models will involve reduced engineering changes late in the project, fewer rejected parts, easier installation into the aircraft, and less engineering time required throughout the project.A military program produced SLA models of a duct manifold and air delivery plenum interfaces. The manifold model was used to ensure conformity to primary aircraft structure before fabrication of the high cost, long lead-time item. The SLA plenum interfaces were made to be assembled into a full-functioning mock up in the systems integration laboratory. The production plenum interfaces were critical items for a mock-up that was needed to meet contract schedules, but delivery was delayed by the supplier by several months.

The production plenum interfaces were critical items for a mock-up that was needed to meet contract schedules, but delivery was delayed by the supplier by several months. The SLA plenum interfaces were made in under a week and have been successfully installed into the mock-up that will be powered up and shown to customers at a subsequent design review.4.0 SLA IN FABRICATIONRaytheon-Waco has used rapid prototyping in fabrication in a variety of ways, and has plans to expand current capabilities. The most common direct use of SLA in fabrication is in tooling, although there are some indirect uses of SLA that will aid manufacturers in visualizing and understanding parts to be produced.4.1 CASTINGSLA models have been used as tooling for investment casting. ACES2 (solid) models havebeen used as a master pattern for creating epoxy tooling, which was then used to make eXPendable wax patterns for forming the investment shell.

ACES2 (solid) models havebeen used as a master pattern for creating epoxy tooling, which was then used to make eXPendable wax patterns for forming the investment shell. Quickcast3 (hollow) models have also been used directly as eXPendable patterns in creating investment shells for casting. Both methods have proven to be accurate, reliable, and lower-cost alternatives to machined tooling for low quantity investment casting. Most of the casting applications at Raytheon-Waco involve quantities too low to justify machined tooling eXPense, so the ability to reduce tooling cost through SLA is a significant advantage. It is also estimated that cycle times are reduced by at least 25% compared to the more traditional investment casting process.60ACES models have also been used as tooling for precision sand casting. Accurate, scaled models of parts have been used in place of machined tooling for forming precision sand molds, reducing cost and time for tooling by an order of magnitude.

Accurate, scaled models of parts have been used in place of machined tooling for forming precision sand molds, reducing cost and time for tooling by an order of magnitude. The final castings were high quality, with no noticeable differences due to RP tooling. Two castings for use in wind tunnel testing were produced in under one week from beginning the SLA model to casting delivery, and the pattern is available for making more parts in the future if necessary.4.2 RAPID TOOLING FOR INJECTION MOWINGRaytheon-Waco has begun working with a plastics vendor to produce rapid tooling for lowvolume injection molding. The injection molding vendor's tooling engineers will work with ProlE part models to design the mold and then Raytheon-Waco will build the mold with the SLA-500. Raytheon-Waco has also interacted with tooling vendors to produce rapid tooling with SLS4 technology for higher durability molds.

Raytheon-Waco has also interacted with tooling vendors to produce rapid tooling with SLS4 technology for higher durability molds. Rapid tooling will allow Raytheon-Waco to afford to purchase very low-quantity injection molded parts that will save weight and be cost competitive with machined aluminum parts. It is also expected that rapid tooling will reduce the time to receive injection molded parts by at least three weeks per order.4.3 METAL DEPOSITION MASTER TOOliNGThe fITst significant use of SLA in fabrication was in producing a complex-shaped lavatory sink with unique cross-sections. This sink was too deep and irregular for traditional processes such as stamping or deep-drawing to be reliable. The process that was used involved making a plaster tool from a SLA master pattern, and then depositing a nickel alloy on the plaster tool.

The process that was used involved making a plaster tool from a SLA master pattern, and then depositing a nickel alloy on the plaster tool. It is estimated that this process saved Raytheon-Waco over $140,000 for a single aircraft, in addition to the likelihood that the parts would not have been delivered within the required schedule if traditional processes had been used. This process was so successful that it has been used on other aircraft for similar circumstances, and is planned for other such instances in the future.4.4 COMPOSITE PANEL TOOLING AT RAYTHEON-WACO MANUFACTURING FACIliTY The SLA lab is working with the Raytheon-Waco composite panel fabrication personnel todevelop a system for producing tooling. The SLA model will serve as tooling, with the composite panel built directly on the SLA model, or for parts that require autoclaving or heating the SLA model can be a master pattern for producing plaster tooling.

Traditionally, this type of tooling is machined metal, so the potential for saving time and cost is great.4.5 TEMPlATES FOR CABINETS AND TRIMSLA templates can also be used to aid fabrication of VIP aircraft components. The templates will be used by cabinet makers or composite panel producers in forming complicated shapes, and also by machine shops or other vendors that will be fabricating parts to mate with the complicated surfaces. It is expected that templates will reduce the amount of rework required to achieve quality interfaces for complex surfaces that require manual operations to produce.5.0 MODELS USED IN TESTINGThe most significant use of SLA models in testing has been for specialized pressure taps tomeasure duct manifold static pressure while experimenting with various air handling system61configurations. These pressure taps are small tube-like protrusions with a round base that had previously been made out of aluminum at a cost of approximately $85 each.

These pressure taps are small tube-like protrusions with a round base that had previously been made out of aluminum at a cost of approximately $85 each. The SLA pressure taps were produced at a cost of $45 for 100, allowing higher levels of stocking to reduce the likelihood of a program having to wait two weeks or more for machined parts to be made. The SLA pressure taps were designed to be used in testing of air-conditioning and other low temperature applications due to the sensitivity of the SLA resin to heat and humidity. The total program savings from using SLA are $8455 per run.SLA models are also being used in wind-tunnel testing. Scaled models of designs for new under-wing pods for Navy aircraft were produced, and were matched to existing scale aircraft models to test the impacts on flight dynamics from the addition of the pods to the aircraft.

The SLA models were produced in far less time and significantly cheaper than metal castings of the same design, and were able to withstand the stresses of wind tunnel operation. The use of SLA models in the wind tunnel is expected to dramatically increase the flexibility and reduce the cost of wind tunnel testing new designs for aircraft modifications.6.0 SUMMARY OF RAPID PROTOTYPING BENEFITSRaYtheon-Waco has made significant use of the SLA-5oo in the five months it has been onsite. There have been over 300 parts made, with over 150 different designs for at least twelve different aircraft programs. Much of the cost and time savings associated with using SLA models is very hard to reliably quantify, but conservative estimates are that over $350,000 has been saved through finding changes and problems early, reducing fabrication costs, and the reduced cost of in-house SLA modeling versus service bureau costs.

There have also been substantial intangible benefits, such as increased customer involvement early in the design process and improved engineering designs due to the opportunity to visualize designs and create models that can be fitted into assemblies. The largest unquantified benefit from on-site SLA modeling has been increasing the ability of Raytheon-Waco to meet contract schedules through time savings in design, analysis, and manufacturing.7.0 CONCLUSIONRaytheon-Waco has had in-house SLA capability since February, 1996, and has already realized significant benefits from rapid prototyping. These have taken the form of increased customer input early in the design process, improved engineering designs through the ability to better visualize CAD models, and decreased fabrication costs for a variety of parts and manufacturing methods.

These benefits have far exceeded the original expectations in both quantity and variety, and have more than justified the acquisition of in-house rapid prototyping technology. It is expected that rapid prototyping will play an ever-increasing role in Raytheon Waco's efforts to reduce operating costs, compress delivery schedules, and enter the next century as one of the world's premier aircraft modification centers.FOOTNOTES: 1. ProlEngineer is a registered trademark of Parametric Technologies, Corp. (PTC) 2. ACES is a trademark of 3D Systems 3. Quickcast is a trademark of 3D Systems 4. SLS is Selective Laser Sintering, a registered trademark of DTM Corp.62.

s Phase SFF Control System for Silicon Nitride Deposition by SALD/SALDVI by Shay Harrison, Chris F. Costa, Kevin 1. Jakubenas, James E. Crocker, and Harris L. Marcus Institute ofMaterials Science University of Connecticut, U-Box 136 Storrs, CT 06269-3136Abstract: A closed-loop laser scanning and temperature control system has been developed for SALD/SALDVI. Temperature control in SALD/SALDVI because temperature plays a defining role in both composition and deposition rate. The control system for time temperature SALD/SALDVI is presented which provides .STL file interpretation, real control, and laser response modeling, all on a PC. This control system was utilized with the SALD/SALDVI techniques for depositing silicon nitride.

This control system was utilized with the SALD/SALDVI techniques for depositing silicon nitride. Characteristics of ShN4 fabricated shapes are discussed, including composition, morphology, and electrical properties.is especially importantIntroductionSelective Area Laser Deposition(SALD)/SALD Vapor Infiltration(SALDVI) are gas phase SFF approaches to forming a vast array of materials from a laser- induced, localized gas phase reaction1. The deposition of ceramic shapes, such as silicon nitride, without post processing presents one of the promising advantages of gas-phase SFF. Laser-CVD growth of silicon nitride fibers from silane(SiHt) and ammonia(NH3) gas mixtures has previously been shown2. Mixed gas environments of tetramethylsilane(TMS, Si(CH3)4) and ammonia have also experimentally produced silicon nitride deposits.

Mixed gas environments of tetramethylsilane(TMS, Si(CH3)4) and ammonia have also experimentally produced silicon nitride deposits. A more thorough examination of the TMS and NH3 deposition system, using thermodYnamic modeling, was undertaken to understand the temperature and gas partial pressures necessary for formation of silicon nitride. With this including closed-loop control of knowledge, a gas-phase SFF computer control system, temperature, was implemented to verifY the modeling predictions. These experiments were performed with a 150 watt cw Nd:YAG laser(A=1.064 nm), an optical pryometer for temperature control, and an 8 inch diameter stainless steel vacuum chamber(the system is more thoroughly described in reference 3). In addition, the electrical and morphological nature of the formed silicon nitride shapes was examined.Gas-phase SFF Computer Control SystemMonitoring and control of both position and temperature are performed on a PC program.

This software is currently used on 2 gas-phase SFFwritten in-house in Visual Basic language4 systems at UCONN(references 5 to 10 detail the theory on which the control system is based).The motion control can handle single two-dimensional layers or multiple 2D layers taken from a three-dimensional .STL or CAD models. The. STL/CAD capability requires that the 3D solid model be sectioned into 2D layers. Each layer must then be 'scanned' with the appropriate laser scan spacing to create a coordinate file for laser scan control. An in-house translation program takes the coordinate file for each layer and creates motion control code for scanning the laser.

An in-house translation program takes the coordinate file for each layer and creates motion control code for scanning the laser. When creating a SALDVI structure, the control program also accesses a powder delivery program11.The original temperature control design utilized a proportional gain to vary the laser power and consequently adjust the reaction temperature to the desired target temperature, based Plaser=P1aser + K * AT , where P1aser is the laser power, K is the gain, on the following equation:241and t1T is the difference between the measured and target temperatures. This simplified control scheme was inadequate, leading to either large temperature overshoots or slow response times. As a result, a Pill(~roportional Integral Derivative) filter was programmed into the control software.

As a result, a Pill(~roportional Integral Derivative) filter was programmed into the control software. The PID filter functions according to the flowchart in Figure 1.R(t) +E(t)U(t)Y(t)._9J-----t>I.ltID Filter1,__.....I SALD - L DepositionFigure 1: Pill FlowchartR(t) is the desired target temperature, Y(t) is the measured target temperature by the optical pyrometer, E(t) is the error between measured and target temperatures, and U(t) is the Pill In a discrete time domain, where k represents a adjustment signal to the laser controller. equation: the based temperature measurement U(k) =U(k -1) + Kp[E(k) - E(k -1)] + KiE(k) + Kd[E(k) - 2E(k - 1) + E(k - 2)].sampling,the PillsignalisonDue to the non-linearity of the laser power response to the analog control voltage from a Digital-to-Analog Card(DAC), a signal conditioner was employed to apply the PID output signal directly to the laser power.

This conditioner is based on a laser calibration of the laser output power to the DAC input power, according to the equation: V(t) =mU(t) +b, where m is the slope and b is the y-intercept ofthe laser power versus DAC curve.With the Pill in place, the next step was to determine the appropriate Pill parameters for small temperature overshoot and a quick response time. Experimentally, a' trial and error approach could take a considerable amount of time. A first order estimation of the necessary Pill The reaction zone values was obtained using a simple system model on a spreadsheet.

A first order estimation of the necessary Pill The reaction zone values was obtained using a simple system model on a spreadsheet. temperature versus time response at constant laser power during deposition can be modeled using _!- T(t) = P x K(I-e t1) where P is the laser power, K = Pthe following equation:t1T(oo)t1T((0) is the difference between the steady-state temperature and the initial temperature, and11 =1 :: o)J (t15% is the time to 75% of T(oo».( In 1- .J.\ 75Yo PxKIn order to use this modelingspreadsheet, a simple constant power, single line scan deposition experiment must be performed to obtain real values for T(oo), P, and t15%.

After inputting these values into the worksheet, PID settings can be tested and a time-temperature response graph is generated, based on consecutive iterations whereY(k) +a1Y(k -1) =b1U(k -1),ofthefollowingequation:b, =~1- exp(- ::J] , 0, =- exp( - ::J' and to is the sampling time_ The graph details theexpected system response. of temperature overshoot and time to reach the steady-state temperature. This modeling program provided initial PID values of Kp=.OI, Ki=.OOI, and :Kt=.Ol242,for the Nd: YAG laser. With further experimentation, the optimal temperature response and control for the YAG laser used PID parameters that are a factor of 10 less than the modeled values.Thermodynamic Modeling of the TMSINH3 Deposition SystemA thermodynamic consideration of the tetramethylsilane and ammonia gas deposition system was performed using a modified version of software titled CET8912 . CET89 is a Gibbs free energy minimization program for solids, liquids, and gases.

CET89 is a Gibbs free energy minimization program for solids, liquids, and gases. The modifications made at UCONN to the routine involve the ability to calculate free energies over a specified temperature range. With this program, the deposition materials and phases of these products are projected based on what is favored by thermodynamics to be in equilibrium at selected temperatures, pressures, and molar ratios. Gas partial pressure were used as the independent variable. The effect of overall total pressure changes the temperature of transition from one region to another in Figure 2 shows a the diagram below, but tetramethylsilane/ammonia gas mixture deposition map.the net form of the map remains the same.TMS/NH3 Deposition Map Carbon3000 ~ 2700 { - . ..

. . 2400 1 2100 { 1800 1 1500 t 1200 -Si3N4 + Carbon 900 fl:--__ 600 ~ Si3N4 300 -I=-P----+--~-+--...::::::::S+====+====+===:::::.+:==~I I I I xTMS = 50 TorrSiC + CarbonSiC + Si3N4 + Carbon0.10.20.30.6 0.5 0.4 TMS/(TMS+NH3)0.70.80.9Figure 2: TMS/NH3 Deposition Map from CET89Based on the deposition map, a partial pressure ratio of 1:2 for TMS:NH3 was chosen to demonstrate the ability to manipulate the product composition to form ShN4.ExperimentalSelective Area Laser Deposition(SALD) of silicon nitride rods served as the initial focus of experiments to determine the composition of the formed material, utilizing x-ray diffiaction spectroscopy(XRDS). Implementing the PID control computer program, deposition target temperatures from 10000 C to 13500 C were selected to achieve constant temperature growth conditions.

Implementing the PID control computer program, deposition target temperatures from 10000 C to 13500 C were selected to achieve constant temperature growth conditions. These temperatures were based on the approximate region in the deposition map, at a 1:2 TMS/ NH3 gas ratio, where silicon nitride would be expected to form(the experimental region can be seen on Figure 2, along the vertical line between the x's). The target temperature is actually a temperature parameter, with variations from the actual temperature due to differences in size between the heated reaction zone from the laser and the size of the sampling area of the243pyrometer. Some modeling of the relation between the actual and measured temperatures has been performed but a more complete analysis will be performed.

Some modeling of the relation between the actual and measured temperatures has been performed but a more complete analysis will be performed. A sample XRDS pattern in shown in Figure 3, where sample #37 was a rod grown at 1100° C target temperature.Sarrpletm and Alpha Silicon N1rideX-Ray IlffracIion Specba100140120~100 til c 00 S .500, •~,.. 1 .. .' II lil:~ ;~ ~ all• • , "~ ;~ ~ all• • , ":.,. , ,! ..III I il " "1 t •· ·" ::', ,",1 'r' .. :~: t~ t ;l • 'f, " " ,", . _:. :~ f, :, "~' I. '; A :\ ~ :~ ,; "I. ~y it'" ...".,.. sit¥- ........ tI ~: :~.,..~ ~ !,:~~'6ti ~~, J. ,. l k.'i•~1i "" ~, ....~ ii;~....., ""K... tJ:sflf.7t,!-SarPe37 I /......ftJp,a Si3N4/...,40, .- '"!~ i.......20 o 20 ~ ~ ~ 40 ~ 00.... ~.''''If'lI'l • .....lJ .1..".....1InI.trI....~, .IIIII~......

~.''''If'lI'l • .....lJ .1..".....1InI.trI....~, .IIIII~...... _il.u,..,i'~ffi OOffi ~ ~ 00kge, 211leIaFigure 3 : X-ray Diffraction Spectra for a ShN4 SALD Rod and an Alpha ShN4 Standard (Note: A matched peak has an up and a down arrow associated with it)The x-ray pattern shows a small texturing effect along the 211 plane, at 29 = 38.9°. Other rod experiments showed silicon nitride formation up to 1350° C deposition temperatures, which may imply a deviation from the thermodYnamic modeling prediction for the transition temperature from silicon nitride to silicon carbide formation, depending on the accuracy of the temperature measurements.The electrical resistivity of both SALD silicon nitride rods and SALDVI silicon nitride single layers was measured using a mega-ohm meter. Available literature report silicon nitride resistivity values greater than 1014 ohm-centimeters12.

Available literature report silicon nitride resistivity values greater than 1014 ohm-centimeters12. Resistivity(p) calculations are based on the following equation: p =R*Alt , where R is the resistance measured from an ohm-meter, A is the area of the sample, and t is the thickness of the sample. Both rod and layer samples exhibited resistances greater than 100 Mega-ohms, which was the maximum calibrated value of the ohm meter used. Conservatively using this resistance, the SALD/SALDVI silicon nitride displayed resistivity of greater than 109 ohm-centimeters. An enhanced method for accurately determining the resistance is presently under consideration.Infiltration of deposited silicon nitride into alpha-silicon nitride powder was performed at Scan speeds ranged from 20 the 1:2 TMS/NH3 gas ratio in a square raster scan pattern. microns/second to 50 microns/second.

microns/second to 50 microns/second. A preponderance of dense SALD material, with very little244infiltration, resulted from the faster scan speed condition. At the slower, 20 microns/second scan speed, some infiltration could be seen, but there was still a great deal of SALD material growing Figure 4 shows the powder/SALD growth interface of the 20 off the powder surface. microns/second scan speed sample, with some deposited material(silver-white) infiltrated into the powder(black). A methodical assessment of the appropriate scan speed, deposition temperature parameter, and gas ratios will be undertaken to determine the best conditions for silicon nitride infiltration.SALD Growth Off PowderInfiltration [ Into Powder...1 100micronsPowder SubstrateFigure 4SALDVI ShN4 Cross-Section, Scan Speed 20 microns/secondConclusionsA temperature and laser scanning closed-loop computer program was designed and implemented for use with the SALD/SALDVI processes..

This program utilizes a Pill filter to minimize the temperature overshoot and maximize the temperature response time, in an effort to create constant temperature deposition conditions. Thermodynamic modeling, using CET89 software, predicted regions where silicon nitride would be deposited in a tetremethylsilane and ammonia gas environment. Employing the closed-loop program, the deposition of silicon nitride was demonstrated, based on gas mixture pressures and deposition temperatures from the thermodynamics. Further experimental work will focus on the deposition transition temperature from silicon nitride to silicon carbide, with comparisons to the thermodynamic model.

Further experimental work will focus on the deposition transition temperature from silicon nitride to silicon carbide, with comparisons to the thermodynamic model. Multi-layer SALDVI silicon nitride shapes will be attempted, and the electrical characterization of ShN4 will be improved to better judge the quality ofthe material being deposited.Acknowledgments245The authors wish to acknowledge the support of DARPA/ONR, contract #NOOO 14-96-11299, and ONR, contract #NOOOI4-95-1-0978.References 1. K.l Jakubenas, B.R. Birmingham, S.Harrison, IE. Crocker, M.S. Shaarawi, IV. Tompkins, I Sanchez, & H.L. Marcus, "Recent Advances in SALD and SALDVI", 7th International Conference on Rapid Prototyping, San Francisco, CA, April, 19972. F.r. Wallenberger, & P.C. Nordine, Journal of Materials Research, Volume 9, Number 3, March 1994, pgs. 527-5303. S. Harrison, IE. Crocker, T. Manzur, & H.L.

s Phase Solid Freeform for Fabrication of Three-dimensional Ceramic StructuresClayton Weiss and Harris L. Marcus Institute of Materials Science and CMBE Department University of Connecticut Storrs, CT 06269-3136AbstractSolid free form of ceramic materials can be achieved by deposition from the gas phase.The Selective Area Laser Deposition, or SALD, technique can be utilized to make ceramic depositions with a uniform chemical composition. In order to make all classes of ceramics, including carbides, nitrides, and oxides, selection of a precursor is an essential step. Often the correct precursor for the deposition requires a special environment, namely, one that can be uniformly heated. System design for a heated deposition chamber is discussed as well as preliminary tests of the system functionality.

System design for a heated deposition chamber is discussed as well as preliminary tests of the system functionality. Silicon Carbide depositions were performed as a means of evaluating system parameters.Ceramic materials can be made without a binder or post processing steps when theceramic material is directly deposited into the final net shape. By using Selective Area Laser Deposition, SALD, it is possible to induce ceramic growth by controlling the movement of a laser spot across the surface of a substrate. A gaseous environment, which contains ceramic precursor gasses, will selectively decompose in the area of heat provided by the laser substrate interaction. The combination of focused laser, X-Y optics positioning stage, and vacuum chamber with a controllable environment, all controlled by a computer, create the right conditions necessary for the SALD process.

This can also be applied in the Selective Area Laser Deposition Vapor Infiltration, SALDVI, mode where deposition is into a powder bed. Both deposition systems can be used for the creation of arbitrary three-dimensional layered geometries and by extension three-dimensional fill volumes for joining ceramics.1Carbides and Nitrides have been deposited using this system, in particular Silicon Carbide and Silicon Nitride, with some success.2 Silicon Carbide is used as a test case to demonstrate the efficacy of the deposition system, as all of the constituents of the final deposit, SiC, are found in Tetramethylsilane, which can be easily vaporized due to its high equilibirium vapor pressure. Other gasses are used in the SiC deposition to assist in temperature control and deposition purity. Though Tetramethylsilane is used as simple ceramic precursor, more complex precursors can also be used if the right conditions are met.

Though Tetramethylsilane is used as simple ceramic precursor, more complex precursors can also be used if the right conditions are met. Silicon Nitride for example is a product of the combination of two precursors Tetramethylsilane and Ammonia, where another level of complexity is needed and considerations such as gas ratios must be more readily considered. The system can be used to create oxides as well as nitrides and carbides, as was previously shown.3 The production of oxides using this deposition has a unique set of challenges associated with it.To demonstrate the potential for oxide freeform, deposition of Al2O3 has been explored. 4Precursor selection is a challenge for this oxide. CVD of Al2O3 is often performed with conditions not suitable for the current SALD system, such as temperatures in excess of 500 C and flowing precursor gasses.

CVD of Al2O3 is often performed with conditions not suitable for the current SALD system, such as temperatures in excess of 500 C and flowing precursor gasses. In order to create Al2O3 and other oxides, as well as to improvethe efficiency of depositions in general, it was necessary to design a system that could be operated at an elevated temperature so higher partial pressure of the reactive gases can be used.Certain precursors particularly metal organic precursors, which could provide the neededelemental constituents, are difficult to utilize for a number of reasons. Often, they exist as a powder, in their as received form. Powdered precursors do not usually have a high enough vapor pressure to produce the yield needed for this type of deposition.

Often, they exist as a powder, in their as received form. Powdered precursors do not usually have a high enough vapor pressure to produce the yield needed for this type of deposition. The precursor powders can be heated, but then will display unreliable vaporization characteristics.5 In order to get past this type of problem a liquid precursor could be used but nevertheless the vapor pressure could be low with many organometallic liquids. Trimethylaluminum, is just such a precursor. Trimethylaluminum is used as the Aluminum source for the deposition of Al2O3; It has been used in the process of Atomic Layer Deposition but not frequently as a Laser CVD precursor.6 To use this chemical in the SALD system with the need for higher yields in mind, it was necessary to redesign the deposition system so that operation at elevated temperatures is possible.To Make the SALD system capable of higher temperature operation (ambient to 350 C), several design features were incorporated.

The Chamber was designed to create as uniform an environment as possible. A simple flanged cylinder was chosen as the main reaction area. A series of computer controlled band heaters act to heat this chamber. The temperature of the chamber is controlled by a temperature-PID control loop operating over three distinct thermocouple-heater regions. All incoming gasses feed into a single heated manifold, also PID controlled, and the precursors can be individually heated. Seen below is an image of the deposition chamber in its current iteration.Figure 1: Heated SALD Deposition Chamber: Pictured above are the heated chamber, X-Y positioning system and optical train, and heated precursor and gas manifold.The primary purpose for heating the system is to allow for increased vapor pressure of the reactants and consequently a higher potential deposition yield. Some of the precursors that have been explored will not be viable for this process if they are not heated.

Some of the precursors that have been explored will not be viable for this process if they are not heated. In general, the precursors would be heated to a certain temperature to achieve a desired equilibrium vapor pressure. The gasses would then be introduced to a chamber slightly hotter than the precursors partial pressure equilibrium temperatue., so that condensation of the precursor gasses will be prevented within the deposition chamber. An example of the vapor pressure verses temperature chart of Trimethylaluminium is shown below.Figure 2: Vapor Pressure Curve of Trimethylaluminum7: At ambient the vapor pressure is less than 10 torr which limits the deposition rate. Elevating the system temperature will greatly increase the rate of precursor influx; a vapor pressure of 100 torr at ~ 70C would be suitable for deposition.In a similar manner to above, the vapor pressure of Tetramethylsilane can also be increased.

It is not necessary to raise the temperature of TMS in order to achieve a workable vapor pressure, as the equilibrium pressure is higher than 600 torr at 20 C.8 SiC does however make a good test of the system as it is relatively stable and not pyrophoric as TMA is.There are several parameters that can be varied during the deposition runs. Theseparameters include: gas composition and mixture ratios, laser power, focus and modulation, number of passes per layer of SFF structure, temperature of the chamber, and laser scan speed. The test of the system was performed with the following parameters held constant: laser power, modulation (duty cycle), and gas concentration. The aim of the experiments was to vary several parameters and then to repeat the experiments at a series of elevated temperatures.A laser power of 3.5 watts, CW, generated from a 1070 nm fiber laser, was used with afocused spot of 90 micron. The substrate was 3mm thick sandblasted Titanium.

The substrate was 3mm thick sandblasted Titanium. Reaction Gasses used were, 50 torr TMS, 300 torr Helium, and 50 torr Hydrogen. Scan Speeds of 10, 50, 250, and 1250 microns per second were used in combination with 1, 5, and 10 superimposed scans per (hypothetical) layer. The object of the tests was to vary scan speed and number of passes over the same spot, in order to understand the deposition characteristics on this substrate. This information would then be used for the construction of multiple layered arbitrary shapes of Silicon Carbide. Deposition results are summarized below.Figure 3: Single layer deposition runs of SiC on Titanium with scan speed varying from left to right, 10, 50 and 250 microns/ second. Visible light microscopy was used with a 20X objective.The scan speeds above resulted in very fine deposits, well bonded to the titanium substrate. The scan speed of 50 microns per second resulted in the most even deposition across the beam profile.

The scan speed of 50 microns per second resulted in the most even deposition across the beam profile. The scan speed of 250 microns per second resulted in the thinnest deposition layer above although a speed of 1250 microns per second was also tested and resulted in an uneven and very light deposition not easily imaged by optical microscopy.250 microns per second was chosen as the scan speed to test multiple co-spatial passes of5 and 10 passes per line. In this case the positioning capabilities were also tested. The images below are segments of 2mm radius circles of SiC deposited onto Titanium substrates.Figure 4: Both depositions above were made with a scan speed of 250 microns per second. The left deposition was made with five consecutive passes of the laser and the right deposition was made with 10 consecutive passes.The highest scan speed, of the three viable scan speeds, was chosen because multiple passes at the lowest scan speed resulted in runaway growth.

This runaway growth resulted from excessive heating of a small protrusion, which then grew upward and away from the surface.The test of the heated chamber showed that the chamber could be maintained at a temperature of 50 C with a variation of ±5 degrees over the different temperature zones. Equipment failure prevented the running of high temperature deposition tests in time for this paper.A temperature controlled deposition chamber will facilitate the formation of oxides withlow vapor pressure precursors. Oxides of interest are Al2O3, which requires special oxidizers and a metalorganic source, and hydroxyapatite, which will require two dissimilar precursors, combined to make a complex oxide with an additional outside source of oxygen. Tests of the system with SiC precursors will illustrate the potential for more novel deposits. The system will be used to achieve the freeform of ceramic materials that have not been deposited by selective laser deposition previously.1 C. M.

The system will be used to achieve the freeform of ceramic materials that have not been deposited by selective laser deposition previously.1 C. M. Weiss, M. Aindow, H. Marcus, “Ceramic Joining by Gas Phase Pulsed Laser Processing,” Proceedings of the Solid Freeform Fabrication Symposium, The University of Texas at Austin (2009) 2 J. E. Crocker, H. Wei, L. L. Shaw and H. L. Marcus, “SALDVI of SiC into metal and ceramic powders,” Proceedings of the Solid Freeform Fabrication Symposium, The University of Texas at Austin (2001), pp. 163–169. 3 I. M. Ghayad, E. Geiss, J. E. Crocker and H. L. Marcus, “Spot joining of Si3N4 and SiC ceramics using selective area laser deposition (SALD) technique,” Proceedings of the Solid Freeform Fabrication Symposium, The University of Texas at Austin (2001), pp. 170–174 4 C. M. Weiss, H.L. Marcus, “Al2O3 Precursor Evaluation for SALD Joining," Proceedings of the Solid Freeform Fabrication Symposium, The University of Texas at Austin (2010) 5 H.O.

s-Phase Laser-Induced Pyrolysis of Tapered MicrostructuresGas-Phase Laser-Induced Pyrolysis of TaperedMicrostructuresJames Maxwell Joseph.Pegna Eric Hill Rensselaer Polytechnic Institute Troy, New YorkAbstract:freeform fab Gas-phase Selective Area Laser Deposition (SALD) is a useful rication tool for the prototyping of simple three-dimensional microstruc tures. Using this method, slender graphite and nickel rods of various tetracarbonyl, respectively. diameters were grown from ethylene and nickel By varying the laser power during growth, tapered cone-like structures were also generated. Rod diameters and material morphology were correlated with the SALD process parameters to demonstrate the mechanisms through which steady-state rod growth occurs--and through which it can be controlled.

Rods, and other similar microstructures, have many useful applications, and are preliminary building blocks for further modelling and development of the SALD process as a micromachining tool.Keywords: Laser chemical vapor deposition, LCVD, SALD, 3-dimensional growth, micro fabrication, rapid prototyping, micro electro mechanical sys tems, pyroiysis, graphite, nickel, rods, cones.1.Int.roduct.ionhas been explored extensively as a Its primary appli cation has been the direct-writing of conductive traces on custom integrated 1989; McWilliams, circuits and higher-level packaging [Adler, result, 1984]. As on the deposition of scanned metal lines on silicon and quartz substrates, usually with film thick nesses less than 5 microns.

As on the deposition of scanned metal lines on silicon and quartz substrates, usually with film thick nesses less than 5 microns. Both pyrolytic and photolytic LCVD have been stud ied.Overthe past decade,laser-induced chemical vapor deposition (LCVD)thin-film prototyping tool.1990; Garrido, focusesthe available literature laIn contrast,SALD (or 3D-LCVD) structuresis a 3-dimensional mode of pyrolytic LCVD, in remains the in 3D-LCVD the and freewhere much Figure 1. While i.e. same, kinetics of growth cease to be determined by substrate properties, form growth of deposit-on-deposit is predominant.tallercan be generated,such asthe rodsshownthe basic growth mechanism for both processes alaser-induced pyrolysis ofgaseous precursor,1.

For more information on 2-dimensional LCVD and its applications, please refer to one of the following references:(early papers) Allen, 1979; Deutsch, 1979;(Short Reviews/important papers) Bauerle, 1984; Ehrlich, 1984; Bauerle, 1990; Baum, 1992;(Books) Ehrlich, 198919951. Ma-xwell, 1. Pegna, E. Hill, reprinted with permission143Gas-Phase Laser-Induced Pyrolysis of Tapered MicrostructuresInitial Power = 6700 mWSteady-State Power: (Running from Left to Right) 3500 mW, 3750 mW, 4250 mW Chamber Pressure = 550 mbar C2H4 Flowrate = 175 sccm H2 Flowrate = 100 sccmFig. 1: Graphite RodsNote that it is not always possible to grow a 3-dimensional microstructure thin-film LCVD process--even if the dwell-time the deposit will only this is termed self the thermophysical properties of the deposit determine whenusing an otherwise-successful of the scanning laser is increased without bound. Often, grow to a certain thickness, and then growth will cease; limiting.

Often, grow to a certain thickness, and then growth will cease; limiting. usually, such height-limiting occurs. Some self-stop mechanisms are:1.The Optical Self-stop Mechanism: absorbance of the deposit is less than that of the substrate, so that less incident power is absorbed as the film grows, thereby lowering the temperature and subsequent growth rates.

Some examples of this effect can be found in the literature: [Allen, 1979; Meunier, 1992, Hopfe, 1993]2.The Conductivity Self-stop: substrate/deposit exhibit athermal conductiv to laser less conductive subity mismatch such that it becomes progressively more difficult heat the deposit strate.) Examples are:(usually a conductive deposit on a[Allen, 1983]3.The Melting Self-stop:the deposit begins to evaporate/melt under the beam focus, eventually reaching an equilibrium conditioll with the growth pro cess.[Westberg, 1992].is in LCVD processes which are not self-limiting--a property which allows 3D-LCVD to be used as a freeform fabrication tool for micro elec In this paper, we willshoYl. how the precursor partial tro mechanical systems. two dissimilar pressure and incident materials, nickel and carbon. The thermal conduc tivity--about 71.8 w/m K at 1000 K--vs.

how the precursor partial tro mechanical systems. two dissimilar pressure and incident materials, nickel and carbon. The thermal conduc tivity--about 71.8 w/m K at 1000 K--vs. 1.3 W/m K for graphite (.1 to layers).Ourinterestlaser power affectthe 3-D growth of the much higherformer has2.Experiment.alFor our 3D-LCVD experiments, we constructed a system which allows growth of a variety of metals and their carbides. It consists of a reaction chamber, gas delivery system, optics table, and laser.© 1995J. Maxwell, J. Pegna, E. Hill, reprinted with permission144Gas-Phase Laser-Induced Pyrolysis of Tapered MicrostructuresThe 3D-LCVD reactor consists of a custom, SO-cm long, quartz tube with sev eral ports for viewing and laser input. The chamber is connected to a pumping station via a manual gate valve which sets the total system pressure during operation.

The chamber is connected to a pumping station via a manual gate valve which sets the total system pressure during operation. The vacuum chamber and gas-delivery system are enclosed within a and the quartz tube is suspended from a ventilated hood for safety purpgse$, heavy superstructure to isolate the system from vibration.reactive gas is passed thrgugh such as·H2 or a source for binary compound such as ethylene, a powder diffuser where$oTid-phaseprecursors are heated and sublime--or else liquid precursgr. The carrier gas are passed through a bl1bbler to evaporate a and evaporated precursor reaction chamber.

The carrier gas are passed through a bl1bbler to evaporate a and evaporated precursor reaction chamber. The into the subsequently pass ratio and quantity of gases delivered to the chamber are determined by flow meters in the system are con trolled by an MKS 3-channel flow ratio controller.To deliver a precursor tg the vacuum system,a gas, either aand pneumaticswitching valves.Flow metersis made using a high-temperature, differential, capac and the high-vacuum itive manometer, region behind the gate valve. Absolute pressure is measured at tem the lower perature, high-vacuum region by a Baratron manometer and dial gauge, allowing absolute measurement of the chamber pressure indirectly.Pressure measurementattached between the reaction chamberForthe carbon deposition, pyrolysis of ethylene, C2H4, was used accordingto the overall reaction:( 1 )Ethylene was chosen as it adsorbs readily onto silicon and graphite attemperature, 1992].

It has an apparent activation energy of 78-90 kJ/mol.and dissociatesreadily atattainable temperaturesroom [Westberg,i.e. Ni(CO)4' was employed; and has an ambient vapor pressure (1000 mbar) at 43·C. As a result, it is one of the fastest known pre have been peak normal deposition rates cursors reported [Herman, 1983]. Ni(CO)4 dissociates according to the reaction:For nickel deposition,acommon LCVD precursor,tetracarbonyl nickel,this compound melts at2S'C,for LCVD;upto 100~m/sNi(CO)4(g) --> Ni(s) + 4(CO)(g),(2 )with an activation energy of 48-49 kJ/mol. The compound also decomposes at relatively low temperatures, beginning at about 1S0'C, and hence requires lower laser also rather fume hood. To supply the toxic, Ni(CO)4 to the chamber without condensation, we cooled a small sample cylinder of the compound in ice water, least S K below the cleanroom temperature of 20·C.

To supply the toxic, Ni(CO)4 to the chamber without condensation, we cooled a small sample cylinder of the compound in ice water, least S K below the cleanroom temperature of 20·C. Vapor pressures up to 200 mbar were possible with this setup.power densities is flammable,to effect growth. Unfortunately,itisand must be handled within aleaving the vapor source atIn both cases, graphite substrates (Alfa/Aesar #10832) were employed, which were together with a brief vacuum bake to remove water and organic residues. The graphite is a useful substrate thermal conductivity which tends to broaden the ini as it has a high lateral tial to the graphitic layers is also very lOW, inhibiting conduction away from the surface layers. Finally, the substrate surface is non-specular, allowing easy visual ization of the focal spot, and it provides many natural sites for nucleation.rinsed in acetone and isopropyl alcohol,laser-induced growth zone.

The thermal conductivity perpendicular© 19951. Maxwell, J. Pegna, E. Hill, reprinted with permission145Gas-Phase Laser-Induced Pyrolysis of Tapered MicrostructuresThe laser used to induce growth was a Coherent-model argon ion laser, with these a maximum output of 8 W (multi-mode) at wavelengths, pyrolytic graphite and nickel have normal spectral reflectances of approximately 0.7 and 0.6 respectively. Gaussian spot radii of 25 ~m (1/e 2) found that while or greater were used throughout the initiation of growth on the substrate depended greatly on the power den sity--requiring small 3-D, steady-state growth of rods was affected much less by the beam waist. The opti length cal achromat. The theoretical spotsize of the optical system was: 5 ~m, with a max imum power density of 200 mw/~m2. TO attenuate the polarized beam, liquid and beam-splitter are crystal polarization rotator a Macintosh computer utilized; DrA board.the 488/514 primary lines.

2: Carbon Axial Rate and Rod Diameters vs. Power3. Result.sTO map the growth process, we conducted a set of experiments where the pro cess parameters and precursor pressure) were varied while the beam remained stationary, yielding needles of various diameters and heights, such as those shown in Figure 1.(in this case the totallaser power© 19951. Maxwell. 1. Pegna, E. Hill, reprinted with permission146Gas-Phase Laser-Induced Pyrolysis of Tapered Microstructures3.1 Carbon Growthto characterize the growth of a moderate conductor the case of carbon on on a substrate of similar thermal conductivity; laser graphite. Preliminary results for powers and pressures are given in Figure 2. Note. that the axial growth rate of the needles approached an asymptote of about 4 ~m/s as the laser power rose to values greater than 4000 mW, limiting along the indicative of mass transport along these lines will be conducted to support centerline.

Further this claim conclusively.Our first objective wasthis wasasequence ofrods grown at constantstudiesrising from a threshold power of about 2750 mw to a maximum diameter of about 270 ~m at pow ers above 6500 mW.Somediffusion limiting of the radial growth rate appears to the total chamber occur beginning at laser p.owers above 5500 mW. pressure was maintained at 550 mbar, while the C2H4 and H2 mass flowrates were 175 sccm and 100 sccm, respectively.The steady,..state rod diameter also varies with laser power,In all cases,InitialeXPeriments where the chamber pressure was altered during growth the steady-state diameter to indicate the radial growth was primarily determined by surface temperature and the thresh(between 150 and 550 mbar) did not appear of initiate at pressures below 150 mbar. The constant needle diamet19rs that reaction kinetics.

The constant needle diamet19rs that reaction kinetics. However, we do not yet have an explanation for old-pressure effect.to affectthe rods greatly,although ne19dle growth from ethylene was difficultLaser Power[mW]8000'1000".60005000"""'-. /40003000~""'-. // '200U1000100200300400500Time [s]Pressure=532 mbarC2H4 Flowrate=175 sccmH2 Flowrate=100 sccmFig. 3: Tapered Graphite Structurestructures of also attempted to grow several We in shown is power; the graphite Figure 3. Note the inset graph where the total time is given. One would expect an hourglass shape from the power profile, but the lower hour glass bulb was much smaller than the upper half. This is tentatively explained by the greater initial conduction to the substrate during the commencement oftapered axi-symmetric such micro-shape one laser power vs.byvaryinglaser© 1995J. Maxwell. J. Pegna. E. Hill.

Maxwell. J. Pegna. E. Hill. reprinted with permission147Gas-Phase Laser-Induced Pyrolysis of Tapered Microstructuresgrowth. Data of rod diameter vs. tapered needle corresponds closely to the straight plot above.laser power measured on the upper bulb of the rod diameters given in the3.2 Nickel DepositionGrowth of nickel microstructures proved to be more challenging than the facilitating rapid carbon growth. While the decomposition temperature is low, growth, allowing solid carbon particulates to form at an amazing rate. Rapid convection past the horizontal needles was observed during growth, and the carbon particulates accumulated quickly on the substrate, often obliterating our samples.the CO reaction by-productstended to dissociate as well,Nickel Needle (2750 mw)Cratered Rods(>8000 mw)-->Fig. 4: Nickel MiicrostnlctiUJ"eSthan simple rods.

4: Nickel MiicrostnlctiUJ"eSthan simple rods. This can be most limitations induced by likely explained by transport the rapid efflux of reaction products--four CO groups for every Ni atom depos ited. Black smoke was often ejected from the focal site in bursts. The cause of smooth polycrystalline lower the periodic bursts is unknown. At rods could be grown, such as that shown in Figure 4. The apparent grain size in the smooth needles was approximately 5 microns, the size rising rapidly above incident powers of 5000 mW.At highlaser powers,crater-like structuresformedratherlaser powers,© 1995 J. Ma>;well, J. Pegna, E. Hill, reprinted with pennis"ion148Gas-Phase Laser-Induced Pyrolysis of Tapered MicrostructuresSeveralimportant observations can be made aboutthe growth of nickel vs. than their graphite thermal conductivity--a property which the same the nickel needles onto the substrate stnall attachment decomposition temperature of Ni(CO)4graphite.

than their graphite thermal conductivity--a property which the same the nickel needles onto the substrate stnall attachment decomposition temperature of Ni(CO)4graphite. First, counterparts, owning to nickel's higher broadens the heated surface area over which deposition may occur. For reason, was very large, while zone. the allows nucleation over a comparatively broad area.the nickel needles were much broaderthe initial attachment point ofrods nucleated from aIn addition,400Steady-State Rod Diameter [pm]350300250200/ v/~-------:.,.--~15010050 o2.533.54.5 Laser Power [W]456Steady-State Axial Growth Rate [Jlm/s]5.554.5//'-/~.------~4......../3.52.533.544.55Laser Power [W]Fig. 5: Nickel Axial Rate and Rod Diameters vs. Power(static fill) were employed throughout the experiment, and rods could be initiated at pressures less than 25 mbar.

5: Nickel Axial Rate and Rod Diameters vs. Power(static fill) were employed throughout the experiment, and rods could be initiated at pressures less than 25 mbar. The axial ethylene-grown structures, than that of growth the the highest powers. The axial rate and steady-state diam exceeding 18 pmfs at eter for several smooth nickel rods can be seen in Figure 5. Further data will be necessary to determine the onset of diffusion limitation and to conclu sively characterize the process for shape-~orming.Ni(CO)4 vapor pressures of 200 mbarrate was higherthe deposits was determined using energy dispersive X-ray indicate that spectroscopy (BDS ). Results is the percentage of carbon, oxygen, and tungsten contamination in the deposit small, in the sample because W(CO)6 has been previously evaporated in the system. The carbon and oxygen were left over by-products.

s-Phase Selective Area Laser Deposition(SALD) Joining of SiC Tubes with SiC Filler Material by S. Harrison and H.L. Marcus Institute of Materials Science University of Connecticut 97 North Eagleville Road Storrs, CT 06269-3136 Phone: 860-486-4410 Fax: 860-486-4745 Abstractjoining together ceramic components with ceramic filler material, known as Selective Area Laser Deposition(SALD) Joining, was utilized in fabricating joined silicon carbide structures. Specifically, silicon carbide tubes were 'welded' together by depositing silicon carbide from a gas phase reaction. Two different precursor environments were examined, one a tetramethylsilane/hydrogen mixture and the other composed of methyltrichlorosilane. The quality of the joints were examined by In addition, the composition and morphology of the bend tests and hermeticity measurements.

The quality of the joints were examined by In addition, the composition and morphology of the bend tests and hermeticity measurements. silicon carbide deposit was studied and is discussed here.laser-driven, gas-phase based SFF technique forIntroductionThe technology for fabrication of robust ceramic parts lacks a proven and reliable method for assembling together complex designs from simple shapes. The challenges presented in joining together ceramic shapes arise from the very properties that make ceramics attractive. Their high melting temperatures preclude traditional welding techniques that utilize base In fact, silicon carbide and silicon nitride sublimate from the solid phase to material melting. their constituent gases and therefore display no melting characteristics. The poor wettability of In attacking ceramic surfaces with most metals presents many difficulties in brazing efforts.

their constituent gases and therefore display no melting characteristics. The poor wettability of In attacking ceramic surfaces with most metals presents many difficulties in brazing efforts. these challenges, a variety of novel ceramic joining procedures have been devised and examined. Diffusion bonding forms ceramic joints by diffusion of atoms across the joint seam at very high processing temperatures. Analogous to solid-state sintering, diffusion bonding suffers from the characteristic low diffusivity of ceramic chemical species and low ductility. Efforts to avoid these shortcomings have utilized metal foils, such as nickel, nickel-chromium and aluminum boride, as interlayers at the joint seam to enhance the diffusion and ductilityl,2,3.

Efforts to avoid these shortcomings have utilized metal foils, such as nickel, nickel-chromium and aluminum boride, as interlayers at the joint seam to enhance the diffusion and ductilityl,2,3. Reactive metal brazing operates by using a metal filler in a traditional braze mixture to react with the ceramic surface and produce a more stable, wettable surface material4 • The most promising brazes are silver-copper alloys with titanium or zirconium mixed in as the reactive agents5 ,6.This research program applied the gas-phase, laser-driven Solid Freeform Fabrication technique of Selective Area Laser Deposition(SALD) to the ceramic joining problem, in a process called SALD Joining. The initial, proof-of-concept investigation of this process at the University of Texas showed its feasibility by linking together alumina tubes, silicon carbide tubes, and silicon nitride tubes7 ,8.

The initial, proof-of-concept investigation of this process at the University of Texas showed its feasibility by linking together alumina tubes, silicon carbide tubes, and silicon nitride tubes7 ,8. The UCONN research efforts studied in-depth the attachment of silicon carbide tubes together using silicon carbide filler material deposited in the SAW Joining process. The ultimate goal is to show the ability of gas-phase SFF techniques to fabricate separate ceramic parts and then in-situ 'weld' them together to form a complex ceramic structure. The tube geometry utilized in this program was chosen for convenience with the expectation that the SALD Joining procedure will work for other geometries and types of ceramic material, including those fabricated by SFF.537Experimental ApproachSALD employs a high-powered laser beam to induce a thermal decomposition reaction of specific gases inside a vacuum chamber.

The decomposition reaction leads to a desired solid product deposited inside the laser spot heated zone. The laser beam can be scanned by a motion control computer system to selectively deposit material in a desired pattern. SALD Joining utilizes the SALD technique to deposit solid product and use it as a finer material to join together ceramic parts. The UCONN joining chamber contains a rotational device to which the experimental tubes were attached. The joint seam could be scanned relative to the laser in several patterns, from a wiggle pattern back and forth across the seam to a simple stationary beam centered on the joint. A schematic of the tube rotation device is found in Figure 1. This scanning, as wen as the tube rotation speed, was computer controlled by the Microsoft Visual Basic-based SFF Control System software. A stabilized, continuous wave(cw) 50 watt Nd:YAG laser from ExcellQuantronix served as the primary laser, operating at a 1.06 micron wavelength.

A stabilized, continuous wave(cw) 50 watt Nd:YAG laser from ExcellQuantronix served as the primary laser, operating at a 1.06 micron wavelength. The primary gas precursors used were a mixture of tetramethylsilane(TMS, Si(CH3)4) and include hydrogen. methyltrichlorosilane(MTS, SiCH3Ch) and a combination of silane(Si~)and methane(C~).Alternativegasprecursorsforsiliconcarbidedepositionn II II II II IIFigure 1: Schematic of Tube Rotation Joining Geometry Two different types of silicon carbide tubes were used, a clay-bonded tube from Bolt Technical Ceramic of approximately 85 to 90% density and a DuPont Lanxide Composite tube of HiNicalon fiber in a silicon carbide matrix, referred to hereafter as fiber tube. The clay-bonded tubing came in Y.z inch and 1 inch outer diameter sizes, while the fiber tube's outer diameter measured 3/8 of an inch.

The clay-bonded tubing came in Y.z inch and 1 inch outer diameter sizes, while the fiber tube's outer diameter measured 3/8 of an inch. Experiments were run in two different manners, one with deposition temperature monitoring using an infrared pyrometer and one without the pyrometer. When utilizing the pyrometer, tube samples were weighed before and after deposition in order to calculate a mass deposition rate in correlation with the approximate temperature range. Otherwise, the samples were directly placed in the chamber. After a completed experiment, the sample underwent sonication for approximately 20 minutes to remove as much polymer reaction by-product as possible and then photographed using a Snappy frame-grabber video card and a color CCD camera. Further analysis was performed after the samples were catalogued.5383 display two successfulto hydrogen.~=-=:...-..:..::..

Further analysis was performed after the samples were catalogued.5383 display two successfulto hydrogen.~=-=:...-..:..::.. ESEM Image of Clay-bondedESEM Image Close-up of Tube-JointDeposit InterfaceMechanical testing of end of the joint remained stationary upward jointsjoints was accomplished using a cantilever-type bend test. One a fixture while the other end was forcibly bent by the strength of the severalof a crosshead on an Instron tensile testing machine.showed low strength compared to that of the monolithic tube standards. AA""A."A"""''''''''' quality of several joints was determined by connecting the samples to ...,...,""... and its mechanical vacuum pump by means of a Swagelok fitting and running vacuum pump and the vacuum chamber allowed to equilibrate, thus giving a measure of the lowest possible 1, along with the vacuum leveljoined structures were exposed topossible for a monolithicjoints.

The results appearstandard.539Microhardness evaluation of the SALD material was performed with a Vickers indenter at a 500 gram load and 15 seconds indenter dwell time. The results are displayed in bar graph form in Figure 8. A total of 20 indentations were taken, and the average Vickers hardness number was 1446 kglmm2 . The pure silicon carbide Vickers number has been reported as 33009.Various samples of the silicon carbide SALD material underwent x-ray diffraction spectroscopy(XRDS). All showed the three peak configuration, at d-spacings of 2.52, 1.54 and 1.31 Angstroms, characteristic of beta silicon carbide, as has consistently been reported in silicon carbide CVD literature lO ,ll,12.

Higher temperature deposition material, in the 1300 to 1500 degrees Celsius range as measured by the pyrometer, displayed an additional peak at d = 3.35 Angstroms corresponding to the main carbon/graphite lineFigure 6: ESEM Image of Fiber TubeFigure 7: ESEM Image Close-up ofJointDeposit InterfaceSample Y2" OD Clay-bonded tube standard Y2" OD Clay-bonded tube sample SJ#18 Y2" OD Clay-bonded tube sample SJ# 32Steady..State Vacuum Level(in9 74 116..:::..=~.:::...:.' Hermetic Quality of Tubes1~~:....=.:.Vickers Hardness Determination ofMaterial540Discussionthis investigation, the SALD Joining process produced successful joints with twodifferent types of silicon carbide tubing. The morphology of the deposit on ~ach tube ~urface, the Environmental SEM images above, shows excellent adherence with interfaces. end joint configuration and not a function of the deposit bonding to the tube.

end joint configuration and not a function of the deposit bonding to the tube. alignment leaves a starting crack at the joint that is the length of the tube wall thickness. work will focus on minimizing this crack by beveling the tube ends with an expected improvement the mechanical behavior. The hermetic seals of the tested joint structures were within an order of magnitude of the standard tube value. The better vacuum sealed joint, at Torr equilibrium pressure, possessed a much smoother deposit morphology, while the other tested joint, at 1 Torr, deposited in a rougher columnar shape. We believe that a smoother deposit structure, Le. more laminar than columnar, will produce better sealing by reducing the microhardness values show regions of high purity porosity and therefore the leak path. silicon carbide as well as regions of very soft deposited material.

silicon carbide as well as regions of very soft deposited material. The porosity inherent in the growth columns likely contributes to the response of these soft regions as does co-deposition of carbon in areas where the deposition temperature shot above approximately 1300 degrees Celsius. The hermetic sealing, microhardness uniformity and compositional uniformity will deposition improve with the full temperature can then be maintained in temperature regions where higher purity silicon carbide forms in a more even, layered manner.Integrated low mechanical response of the joined structures can likely be traced to the buttbutta..·nfon ..""implementation of the temperature control system.silicon carbide material contained first was broadening of the main peak at d = 2.52 Angstroms, two interesting features. leading to a grain size analysis. Over several samples, the average grain size was determined to the range from 5 to 40 nanometers.

leading to a grain size analysis. Over several samples, the average grain size was determined to the range from 5 to 40 nanometers. Further grain size analysis with dark field TEM was be performed and yielded an average size of approximately 30 nanometers. second Intlere:)t1nu! feature was examined only after Nuclear Magnetic Resonance(NMR) magic angle spinning characterization of the SALD material. The NMR data showed unexpected profiles compared to a beta silicon carbide powder sample. Closer analysis of the imaging and of the fine structure of the x-ray patterns around the main peak, in conjunction with computer simulations of silicon carbide x-ray spectra13 SALD material is highly faulted. A publication concerning this examination of the SALD silicon carbide material is expected in the nearx-ray diffraction spectra obtained from the,14, lead to the conclusionConclusionsSALD Joining process offers great potential to solve the problems involvedceramic joining.

Monolithic joints, composed of the same material as the base material being 'welded' together, are possible. A dependable ceramic joining procedure could be integrated into a one step. This research shows that silicon carbide joints can be repeatedly fabricated with good mechanical, compositional and hermetic properties. Future work will focus on improving these properties and include tube end beveling, temperature closed-loop control and use of a second, higher power cw Nd:YAG laser. Higher purity silicon carbide tubes will also be joined together.fabrication system to build up large, complex parts from small, simple shapes541Acknowledgments The authors would like to recognize Drs. J.V. Tompkins and B.R. Birmingham forinitial SALD Joining work at UT-Austin, as well as Dr. KJ. Jakubenas for his work on project as a postdoctoral fellow at the University of Connecticut.

J.V. Tompkins and B.R. Birmingham forinitial SALD Joining work at UT-Austin, as well as Dr. KJ. Jakubenas for his work on project as a postdoctoral fellow at the University of Connecticut. This research was Tnnrtl<:ortl part by ONR Grant #NOOOI4-95-1-0978.Footnotes1.2.3.5.6.7.8."Feasibility Study of the Welding of SiC" by T.J. Moore, Journal of the American Ceramic Society, Vol. 68, 1985, pgs. C151-C153 "Joining Silicon Nitride to Itself and to Metals" by S.D. Peteves, G. Ceccone, M. Paulasto,V. Stamos and P. Yvon, Journal of Materials, Vol. 48, January, 1996, pgs. 48-52 & 74-77(52) "Solid-State Bonding of Silicon Nitride Ceramics with Nickel-Chromium Alloy Interlayers" by M. Nakamura and S.D. Peteves, Journal ofthe American Ceramic Society, Vol. 73(#5), 1990, pgs. 1221-1227 "Joining of Ceramics" by R.E. Loehman and A.P. Tomsia, American Ceramic Society Bulletin, Vol.

Nakamura and S.D. Peteves, Journal ofthe American Ceramic Society, Vol. 73(#5), 1990, pgs. 1221-1227 "Joining of Ceramics" by R.E. Loehman and A.P. Tomsia, American Ceramic Society Bulletin, Vol. 67(#2), 1988,pgs.375-380(377) "The Influence of Brazing Conditions on Joint Strength in Al20Y'Al20 3 Bonding" by H. Hongqi, 1 Zhihao and W. Xiaotian, Journal of Materials Science, Vol. 29, October 1st, 1994, pgs. 5041-5046 "Ag Effect on Microstructures and Strength of SiC/SiC Joint Brazed with Cu-Ag-Ti Filler Metals" by T. Tamia and M. Naka, Journal ofMaterials Science Letters, Vol. 15(#12), June, 1996, pgs. 1028-1029 "Ceramic Joining by Selective Beam Deposition" by lV. Tompkins, B.R. Birmingham and H.L. Marcus, Solid Freeform Fabrication Symposium, Austin, Texas, August, 1995, pgs. 409-416 Joining Ceramics by Selective Beam Deposition by lV. Tompkins, PhD Dissertation, University of Texas at Austin, May, 1998,pgs. 103-1229. Modern Ceramic Engineering, 2nd Edition by D.W.

409-416 Joining Ceramics by Selective Beam Deposition by lV. Tompkins, PhD Dissertation, University of Texas at Austin, May, 1998,pgs. 103-1229. Modern Ceramic Engineering, 2nd Edition by D.W. Richerson, Marcel Dekker Publishing, New York, 1992, pg. 17910. "The Structure of Chemical Vapor Deposition Silicon Carbide" by J. Chin, P.K. Gantzel and R.G. Hudson, Thin Solid Films, Vol. 40, 1977, pgs. 57-7211. "Growth Characteristics of Beta-Silicon Carbide by Chemical Vapor Deposition" by C-H. Chu, Y-M Lu and , 1992, pgs. 3883-388812. "A Morphological and Structural Study of SiC Layers Obtained by LPCVD Using Tetramethylsilane" by A. Figueras, S. Garelik, R. Rodriguez-Clemente, B. Armas, C. Combescure and C Dupuy, Journal of Crystal Growth, Vol. 110, 1991, pgs. 528-54213. "Effect of Stacking Faults on the X-ray Diffraction Profiles of Beta-SiC Powders" by V.V. Cawley, Journal of the American Ceramic Society, Vol. 78(#3), March, 1995, pgs. 774-782 and J.D.14.

"Effect of Stacking Faults on the X-ray Diffraction Profiles of Beta-SiC Powders" by V.V. Cawley, Journal of the American Ceramic Society, Vol. 78(#3), March, 1995, pgs. 774-782 and J.D.14. "Computer Simulations of Diffraction Effects due to Stacking Faults in Beta-SiC:I Simulation Results" by V.V. Pujar and J.D. Cawley, Journal of the American Ceramic Society, Vol. 80(#7), July, 1997, pgs. 1653 1662542.

S-PHASE SOLID FREEFORM FABRICATION OF SiC CERMETS USING SALDVIJames E. Crocker, Leon L. Shaw, and Harris L. Marcus Institute of Materials Science, Department of Metallurgy and Materials Engineering University of Connecticut, Storrs, CT 06268AbstractSelective Area Laser Deposition Vapor Infiltration (SALDVI) is an experimental solidfreeform fabrication (SFF) technique aimed at the direct fabrication of ceramic and ceramic/metal structures and composites. SALDVI uses a layer-by-layer approach in which powders are infiltrated with solid material deposited from gas precursors by chemical vapor deposition (CVD) using laser heating. Experiments have been performed with CO2 and Nd:YAG lasers using the silicon carbide forming gas precursor Si(CH3)4 and Cu, Mo, and Ni metal powders. The microstructure of the resulting SiC/metal cermets was investigated in relation to the processing history.

The microstructure of the resulting SiC/metal cermets was investigated in relation to the processing history. In some cases, the formation of intermetallic silicide phases was observed.IntroductionSALDVI is a layer-by-layer approach to SFF in which porous layers of loose powders are densified by depositing solid material from gas precursors into the pore spaces using laser chemical vapor infiltration (CVI)1. Because the CVI process can deposit a variety of desirable materials in pure form, SALDVI is capable of building ceramic and composite net shapes and shapes with functionally graded compositions such as embedded devices2 in a single machine. SALDVI using ceramic powders and the SiC-forming gas Si(CH3)4 has been previously investigated3. The laser induced processing temperature distribution, gas precursor pressure, laser scanning speed, and the particle size of the starting powder were found to affect the infiltration and densification rates of the bulk SALDVI shapes.

The nanocrystalline nature of the solid material deposited from the Si(CH3)4 gas precursor by laser CVD on solid ceramic substrates has also been well characterized on the micro and atomic scales4. In all of this previous work, the ceramic substrate remains chemically compatible with the vapor deposited SiC throughout the range of processing temperatures experienced during fabrication. In this study, the fabrication of SALDVI cermets using Cu, Ni, and Mo metal powders and Si(CH3)4 gas is explored. The affinity of these metals for forming a range of intermetallic compounds with free Si combined with their lower melting temperatures, more so for the Cu and Ni than the Mo, makes their chemical stability with the vapor deposited SiC uncertain. In the following, the processing temperature and time at temperature are varied to examine the effect on the extent of reaction between the starting powders and the vapor infiltrated matrix.

In the following, the processing temperature and time at temperature are varied to examine the effect on the extent of reaction between the starting powders and the vapor infiltrated matrix. Scanning Electron Microscopy (SEM) with Energy Dispersive X-ray Spectroscopy (EDS) is used to probe the local elemental compositions at each processing condition. X-ray diffraction is used to detect the formation of intermetallic phases in the bulk samples.ExperimentThe SALDVI workstation consists of a vacuum chamber, a powder delivery system, a 50 watt continuous wave CO2 (10.6 m m wavelength) or a 150 watt continuous wave Nd:YAG (1.06168m m wavelength) laser beam, an xy table with scanning mirrors, and an optical pyrometer temperature probe. All components are computer controlled.

All components are computer controlled. In these experiments the surface temperature measured by the optical pyrometer was used in a feedback loop to adjust the laser output power as necessary to maintain a constant surface temperature, termed the target temperature, throughout the experiment. The laser beam spot size was nominally 1 mm in diameter and Gaussian in shape. The Si(CH3)4 gas pressure was 10 Torr. The sample geometry consisted of a single line one beam diameter in width, 20 mm in length, and three layers in height at 120 m m per layer. The Cu, Ni, and Mo powders range in size from a few microns up to 50 m m, are mostly spherical, and pack to an initial powder bed density of about 55 percent. Figure 1 illustrates the size distribution and morphology of the as-received Mo powder. As it was impossible to heat the pure Cu powder to the desired target temperature with either laser, 10 weight percent Si powder was added to improve laser absorption.

As it was impossible to heat the pure Cu powder to the desired target temperature with either laser, 10 weight percent Si powder was added to improve laser absorption. The Ni and Mo powders were processed in the pure, as-received state. Samples were fabricated at target temperatures of 1000, 1100, and 1200(cid:176) C, and at laser scanning speeds of 1.25, 2.5, and 5.0 m m/s. Considering the laserspot size of 1 mm, these scanning speeds result in laser dwell times or average heating times of 800, 400, and 200 s, respectively. Polished cross-sections of each condition were examined in the SEM and the local chemical composition of the Si and Cu, Ni, or Mo was probed with EDS. No attempt was made to quantify the carbon in the sample with this method. Selected samples were also analyzed with X-ray diffraction. Finally, the mechanical behavior of the three cermets was compared in four-point bending.50 mFigure 1.

Selected samples were also analyzed with X-ray diffraction. Finally, the mechanical behavior of the three cermets was compared in four-point bending.50 mFigure 1. Molybdenum powder particles before SALDVI processing.(a)(b)200 m(c)20 mFigure 2. Optical (a) and SEM photos (b) and (c) of centerline region of Mo powder vapor infiltrated with SiC at 1000 C target temperature and 400 s average heating time.169Results Figure 2a shows a typical as-fabricated single line, three layer, 20 mm long SALDVI sample. The centerline of the scan, where the intensity of the laser beam is highest, consists of a region of dense vapor infiltrated powder. The density decreases to that of the initial powder at the sample edges, which can be easily removed by fine grinding. SEM photos of the dense centerline region are shown in Figure 2b and 2c for Mo powder processed at a target temperature of 1000 C and average heating time of 400 s.

SEM photos of the dense centerline region are shown in Figure 2b and 2c for Mo powder processed at a target temperature of 1000 C and average heating time of 400 s. The dense region is about 400 m m wide and shows the nodular surface appearance common to vapor deposited surfaces.A. Cu and 10 wt. % Si powderSEM images of the SiC vapor infiltrated Cu-10 wt. % Si powder are presented in Figure 3(a-c) for a target temperature of 1000 C and for average heating times of 200, 400, and 800 s, respectively. Figure 4 shows the result for a target temperature of 1200 C and an average heating time of 200 s. The original Cu particles are clearly distinguishable from the surrounding vapor deposited SiC matrix. However, in all cases, EDS measurements of the Cu and Si indicate that the composition within the particles are uniform and are equivalent to that obtained for a Cu5Si standard (8.8 wt. % Si). Small amounts of Cu, up to 5 wt.

% Si). Small amounts of Cu, up to 5 wt. %, were also measured in the SiC matrix within a few microns of the particle surface. The observed compositions showed little dependence on target temperature and heating time. The presence of particle pull-outs in the cross-sections is an indication of the high stresses that exist across the interface between the particles and the matrix due to the thermal expansion mismatch between Cu, SiC, and their reaction products.(a)(b)(c)5 m m5 m m5 m mFigure 3. SEM photos of Cu-10 wt. % Si powder infiltrated with vapor deposited SiC at 1000 C target temperature for (a) 200 s, (b) 400 s, and (b) 800 s average heating times.(a)(b)100 m5 mFigure 4. SEM photo of Cu-10 wt. % Si powder infiltrated with vapor deposited SiC at 1200 C target170temperature for 200 s average heating time.B.

% Si). The gray zone surrounding the particle measures a higher Si content, up to 30 wt. %. In the SiC matrix just adjacent to this gray zone, up to 5 wt. % Ni is detected. Again, similar to the Cu cermet, in Figure 5a there is evidence that some of the reacted particles have debonded from the surrounding matrix.(a)(b)20 m5 mFigure 5. SEM photos of Ni powder infiltrated with vapor deposited SiC at 1000 C target temperature for 800 s average heating time.The effect of the processing target temperature is illustrated in Figure 6. Figure 6a shows a Ni particle from a sample processed at a target temperature of 1000 C for a heating time of 200 s. EDS probes detected no Si at the center of the particle, up to 5 wt. % Si at the edge of the particle, 20 wt. % Si in the zone surrounding the particle, and a few wt. % Ni in the SiC matrix adjacent to the reaction zone.

% Si at the edge of the particle, 20 wt. % Si in the zone surrounding the particle, and a few wt. % Ni in the SiC matrix adjacent to the reaction zone. Figure 6b shows a more pronounced reaction zone from a sample processed at a target temperature of 1200 C for an average heating time of 200 s. The Ni and Si composition in the particle measured by EDS is uniform and richer in Si than that of the Ni2Si phase. The Si content in the regions surrounding the particles is greater than 30 wt. %.(a)5Figure 6. SEM photos of Ni powder infiltrated with vapor deposited SiC for 200 s average heating time and (a) 1000 C and (b) 1200 C target temperature.171C. Mo powderThe Mo powder showed no reaction with the vapor-deposited SiC at a target temperature of 1000 C for the three heating times examined here. No Si was measured by EDS within the Mo powder particles, nor any Mo into the SiC matrix.

No Si was measured by EDS within the Mo powder particles, nor any Mo into the SiC matrix. Figure 7a shows the structure of a Mo/SiC cermet processed at a target temperature of 1000 C and a heating time of 400 s. The lack of pullouts in the cross-section indicates that the Mo particles are well bonded to the SiC matrix. This is not unexpected as Mo and SiC exhibit similar thermal expansion behavior. The solid density, including powder and vapor deposited SiC, is about 95 % in this region. Figure 7b shows the fracture surface across two adjacent layers of a 3 layer Mo/SiC sample. It is apparent that the fracture path goes around, rather than through, the Mo particles. Also, there is continuous vapor deposited material across the layers.(a)20 m(b)20 mFigure 7. SEM images of Mo powder infiltrated with vapor deposited SiC at 1000 C target temperature and 400 s average heating time (a) cross-section and (b) fracture surface.D.

SEM images of Mo powder infiltrated with vapor deposited SiC at 1000 C target temperature and 400 s average heating time (a) cross-section and (b) fracture surface.D. Mechanical BehaviorSALDVI cermet samples processed at 1000 C target temperature and 400 s heating timewere mechanically tested in four point bending. Only one sample of the Ni cermet was sufficiently sound for the test. Three samples each of the Mo and Cu-10 wt. % Si cermets were tested. The failure loads for each are compared in Figure 8. The failure loads for the Mo/SiC cermet correspond to a beam strength of about 50 MPa.1721614121086420NiCu+SiMoFigure 8. Failure loads in 4-point bending for SALDVI samples of Ni, Cu-10 wt.% Si, and Mo powders processed at 1000 C target temperature and 400 s heating time.Discussion The microstructures of the three SALDVI cermets reveal varying degrees of reactionsbetween the metal powders and the vapor deposited SiC matrix.

Both the Cu and Ni powders react with vapor deposited SiC during the SALDVI process at all of the processing temperatures and heating times examined here. For the Cu-10 wt. % Si powder, particles at least 20 m m in diameter reach a uniform composition indicative of a Cu silicide phase even at the shortest heating time of 200 s and lowest processing temperature of 1000 C. Further evidence of Cu silicide formation was observed in preliminary x-ray diffraction measurements at 1000 and 1200 C target temperatures. These results indicate that Si is a fast diffuser in Cu. The observed reactions between Cu and Si are not due only to the presence of the elemental Si in the initial powder mixture however. Similar EDS measurements for a SALDVI cermet fabricated using an initial powder of Cu-50 vol.

Similar EDS measurements for a SALDVI cermet fabricated using an initial powder of Cu-50 vol. % SiC also detected Si within the Cu particles, so at least some of the diffusing Si comes from the vapor deposited material.In the Ni case, the extent of the reactions progressed more notably with increasingprocessing temperature and heating durations. At a target temperature of 1000 C and 200 s heating time, Si diffuses only a few microns into the Ni particle (Figure 6a). Increasing the heating time by a factor of four results in a fully reacted particle up to 20 wt. % Si in the center (Figure 5b). Increasing the processing temperature to 1200 C for a 200 s heating time increased the Si composition measured at the center of the particles higher than 20 wt. % (Figure 6b). The particles sampled here were located in similar regions of each sample, where the temperature and173time at temperature are highest.

% (Figure 6b). The particles sampled here were located in similar regions of each sample, where the temperature and173time at temperature are highest. Because the Gaussian beam shape results in a distribution of temperatures and heating times, certainly other regions in each sample will show varying degrees of reaction. Though the carbon content was not quantified in this study, the reaction between Ni and SiC likely results in free carbon, as Ni does not form a carbide. A diffusion study between Ni and SiC at 1000 C and 1150 C found a reaction zone consisting of alternating layers of Ni2Si and Ni2Si with carbon, consistent with the Ni-Si compositions reported here5. The reaction path from pure SiC to pure Ni at 1150 C in their study suggests that diffusion can occur entirely in the solid state, even though some of the binary Ni-Si silicide phases are liquid below this temperature.

It is not clear if carbon plays a similar stabilizing role in the Cu-Si-C system.For the Mo/SiC cermet, no Mo-Si interdiffusion was detected by EDS for the 1000 Ctarget temperature. This is not surprising as the melting temperature of Mo is 2610 C, so diffusion would be expected to occur slowly at 1000 C. The Mo/SiC cermet shows the highest bending strength of the three cermets. This behavior is attributed to the close thermal expansion match between the two materials resulting in few particle debonds compared to the Cu and Ni cases.ConclusionsThe fabrication of SiC cermets by SALDVI using Cu, Ni, and Mo powders wasinvestigated. The Cu and Ni powders reacted with the vapor deposited SiC at the normal SALDVI processing temperatures and scan speeds to form microstructures with various interesting compositional features, including the formation of intermetallic phases.

No such reactions were observed for the Mo/SiC cermet, although only the lowest processing temperature of 1000 C was currently investigated. The highest bend strength of 50 MPa was measured for the Mo/SiC cermet.AcknowledgementThe authors acknowledge the support for this research by the Office of Naval Research(grant #N00014-95-1-0978).References1. B.R. Birmingham and H.L. Marcus, “Silicon Carbide Shapes by Selective Area Laser Deposition Vapor Infiltration,” Proceedings of the Solid Freeform Fabrication Symposium, The University of Texas at Austin, 1994, pp. 348-355.2. J.E. Crocker, S. Harrison, L. Sun, L.L. Shaw, and H.L. Marcus, "Using SALDVI and SALD with Multi-Material Structures," JOM, 50, 21-23, December, 1998.3. J.E. Crocker, L. Sun, H. Ansquer, L.L. Shaw, and H.L. Marcus, "Processing and Characterization of SALDVI Ceramic Structures," Proceedings of the Solid Freeform Fabrication Symposium, The University of Texas at Austin, 1999, pp. 495-501.4. J.E.

e GE Aircraft Engine Bracket Challenge: An Experiment in Crowdsourcing for Mechanical Design ConceptsCarter, W.T., Erno, D.J. GE Global Research, Niskayuna, NYAbbott, D.H., Bruck, C.E., Wilson, G.H., Wolfe, J.B. GE Aviation, Cincinnati, OHFinkhousen, D.M., Tepper, A. GE Corporate, Fairfield, CTStevens, R.G. GrabCAD, Boston, MAAbstractAn emerging international engineering design trend has resulted from widespread use of social media: a large number of people are engaged in collaborative engineering design activities to build their design expertise through interaction with other designers, to compete for recognition or prizes, or simply for the enjoyment of doing so. The term “crowdsourcing” was introduced in 2005 and implies soliciting contributions from a large group of people, usually an online community, in order to get a broad perspective from various points of view.

This community-generated creativity is contrary to conventional practice in most manufacturing companies, which prefer tight control of engineering designs and practices because they represent key intellectual property and know-how. Recognizing that crowdsourcing represents a potential resource, GE embarked on an experiment to see how a for-profit company might benefit from soliciting new design approaches from this non-traditional source. The design of a specific part, an aircraft engine bracket, was released to an online community of engineers with an invitation to submit improved designs in an open competition. Entrants were encouraged to consider additive manufacturing as the fabrication method. Hundreds of designers submitted concepts and some achieved 80% reduction in weight.Introduction: Open InnovationOpen innovation is the formal practice of engaging the world to help drive innovation, add value, and solve problems creatively and rapidly.

Examples of open innovation include innovation contests, crowd-sourced product development, and innovation networks.i Open innovation is contrary to conventional practice in most manufacturing companies, which prefer to control engineering designs and design practices because they represent key intellectual property and know-how. The term “crowdsourcing” was introduced in 2005 and implies soliciting contributions from a large group of people, usually the online communityii, in order to get a broad perspective from various points of view.Recognizing that crowdsourcing represents a potential resource that can benefit the company, GE embarked on an experiment to see how a for-profit company might engage the open community. Two design challenges were selected, one for GE Healthcare, and one for GE Aviation.

Two design challenges were selected, one for GE Healthcare, and one for GE Aviation. The GE Healthcare focused on tungsten materials for CT detectors.i For the GE Aviation challenge, the topic of this paper, the choice was made to release the design of a rather simple though important part of the aircraft engine to the open community with the challenge to improve its design by reducing its weight. The part is an aircraft engine bracket, and its function is to support the weight of the cowling during engine service – it plays no active role during the operation of the engine. The part, shown in Figure 1, is made of titanium and weighs about 1.8 kg (4 lb.). It was recognized that the part could meet all functional requirements at a lower weight.A cash prize was offered for the satisfactory design of the least weight. A manufacturing technique was not stipulated, though contestants were encouraged to take advantage of modern additive manufacturing methods if they desired.

A manufacturing technique was not stipulated, though contestants were encouraged to take advantage of modern additive manufacturing methods if they desired. Everyone in the world was invited to compete, and almost 700 people chose to do so.Figure 1. The original bracket design. This part was machined from titanium bar stock and weighed 4 lb.The Value of Weight Removed from an Aircraft EngineWeight minimization is often a key engineering design goal: an efficient design is usually a light design that is capable of handling whatever loads are applied without fracturing or permanently deforming under load. Weight is of importance in static applications, but grows in significance in dynamic or moving applications (such as automobiles) and becomes very significant with lifting applications (such as air travel and space flight).A very large number (93,000) of commercial aircraft fly every day worldwide.

Modern aircraft are, of course, heavy vehicles: the Boeing 737 weighs approximately 65 metric tons at takeoff. Recognizing that only passengers and cargo are revenue-generating, airlines are faced with the fact that transporting the airframe, fuel, and engines represents a necessary but unwanted cost of operation. With that in mind, an intriguing question: How much fuel would be saved annually by reducing the weight of each worldwide aircraft by one pound? We will answer the question from two perspectives, that of a businessman and that of a physicist.The CEO’s AnswerIn 2012, a well-known airline flew 109,346,509 revenue passengers on a total of 1,361,558 trips at a cost of $6.12B in fueliii. These values indicate an average fuel cost of $4500 per trip, or $56 per passenger. If we assume that the average passenger weighs 200-lb, including luggage, then the fuel cost is $0.28 per pound per trip.

If we assume that the average passenger weighs 200-lb, including luggage, then the fuel cost is $0.28 per pound per trip. Thus, if one pound could be stripped from all of its aircraft, the airline would save just over $380,000 per year. Scaling this number to include all aircraft worldwide, the savings increases to more than $9.5M per year (> 3M gallons).The Physicist’s AnswerLet’s assume an efficiency, as given by the Brayton Cycle using a compression ratio of 30:1, and assumed losses due to friction, cabin services, cooling, electronics and controls, etc. The energy required to elevate a pound-mass to a height of 10-km (~35,000 ft) is then:However, the weight of the aircraft and fuel must be included in the calculation along with the weight of the payload; perhaps two pounds must be added for each pound of payload, tripling this number.

In addition, fuel cost for taxiing, idle, cruise, approach and landing has been ignored; based on a cursory review of the literature, we will assume that takeoff amounts to 10% of the total flight fuel usage. Further, assuming a jet fuel energy density of 34.7 MJ/L,iv and a cost of $0.8/L ($3.00/gal), the cost per pound is $0.15 per pound per trip. Scaling this number to include all airlines worldwide gives a total savings of over $5M per year (1.7M gallons).The two answers are surprisingly close given the difference is problem solving method and assumptions. It is worth noting that, since most airframes require two, three, or four engines,the incentive for weight reduction in aircraft engines is proportionately higher. We conclude that the incentive for weight reduction in aircraft engines is enormously high.The ContestRecognizing that low weight is a key requirement, GE designers reviewed several parts and eventually selected the bracket of Figure 1.

GE decided to engage GrabCADv, an established community of over 1M engineers to handle the interaction with the global community. The contest was announced in June 2013 and the following requirements were stipulated:Any CAD software could be used as long as a STEP or IGES file was submitted.  The optimized geometry must fit within the original part envelope.  The material was Ti-6Al-4V with an assumed yield strength of 131 ksi at the service temperature (75°F).Minimum material feature size (wall thickness): 0.050 in. Interface 1: 0.75 inch diameter pin. The pin was to be considered infinitely stiff for analysis purposes. Interfaces 2–5: 0.375-24 AS3239-26 machine bolt. Nut face 0.405 in. max ID and 0.558 in. min OD. The bolts were to be considered infinitely stiff.  Loading conditions were stipulated as shown in Figure 2.

Nut face 0.405 in. max ID and 0.558 in. min OD. The bolts were to be considered infinitely stiff.  Loading conditions were stipulated as shown in Figure 2.  The top ten designers would share $20,000 in prizes.The contest generated a great deal of interest and contestants began to upload their entries almost immediately. Recognizing that this was not a typical open design contest where entries would be judged subjectively for their aesthetic appeal, contestants needed clarity on the objective factors of merit to be applied by the judges. Many questions and much online interaction resulted. The contest was closed on Nov. 2013 and almost 700 people had submitted designs. Judging commenced.JudgingJudging proceeded in two phases. The first phase involved selecting the top ten designs with the highest likelihood of meeting the design requirements, and the second involved actually fabricating the top ten and applying the design loads on a test machine.Figure 2 Load Conditions.

Each bracket entry was required to pass finite element analyses to verify that the bracketwould not plastically yield when loaded under the four specific loading conditions above.Phase 1The enclosed volume of all entries was calculated and the entries were ranked by volume from lowest to the highest. Starting with the lightest entry and proceeding until ten leading designs were identified, the following analyses were performed:1. The volume of the entry was confirmed to be totally enclosed within the geometry of the original bracket. Unfortunately, several promising candidates failed this test. The judges acknowledge that some of these entries would likely fit within the constraints of the actual application, and therefore may have worked for the actual application, but were forced to apply the as-stated rules and disqualify these candidates.2. Each entry was meshed with finite elements.

Each entry was meshed with finite elements. Four finite element analyses were performed on each entry, one for each of the load cases. In these analyses, the bolts and mounting surfaces were treated as rigid bodies as described in the rules.3. The maximum von Mises stress was computed within all finite elements. If this value exceeded the yield stress of the material (131 ksi) the entry was disqualified. If the maximum von Mises stress was within 5%, the results were carefully scrutinized to ensure that the finite element discretization did not over-predict these values. In some cases the geometry wasremeshed with a finer mesh for verification, but in none of these cases was the outcome reversed.4. When the lightest ten entries that met the yield stress requirement were identified, analysis offurther designs was stopped.Figure 3 Typical FEA results from phase 1 of the judging.

When the lightest ten entries that met the yield stress requirement were identified, analysis offurther designs was stopped.Figure 3 Typical FEA results from phase 1 of the judging. Von Mises stress is shown, where red indicatesexceeding the yield stress.Phase 2The top ten entries were built from titanium powder in an ARCAM additive manufacturing system. The build direction was consistent for all entries, though it was recognized that some could be built more efficiently in other orientations – this was left for post- contest manufacturing optimization. Support structure was added as deemed necessary by the manufacturing engineer for successful construction; this support structure was removed immediately after the build using hand tools.Several uniaxial tension specimens were built beside the test parts in the AM system to verify that the yield stress given in the rules was correct.

A yield stress of 131 ksi was indeed measured using the 0.2% offset yield stress criterion.A single load fixture, shown in Figure 4 on the left was built to test the brackets under Load Conditions 1, 2, and 3, and the fixture shown on the right hand side of Figure 4 was built to apply the pure torque loading of Load Condition 4. The loading fixtures were placed in a test frame with capability up to 20,000 lb. and each bracket was inserted with new screws to a specified torque. A load-displacement curve was generated and the judges carefully examined the curves to detect any departure from linearity that would indicate permanent deformation of the bracket resulting from plastic deformation. The top three winning entries showed no measurable plastic deformation, though some of the other entries did – these were ranked by a weighted value that included both weight and load at the onset of plastic deformation.Figure 4. Text Fixtures.

Text Fixtures. Test fixtures for Load Cases 1, 2 and 3 (left) and for Load Case 4 (right). The bracket wasfree to slide in slots on the plate for Load Case 4.Figure 5 One of the finalists under test.The Winning DesignsThe top eight designs are shown in the figure below. The winning designer was able to remove 80% of the weight of the bracket without reducing its load-bearing capacity. Additive manufacturing is not a requirement for all the designs, though several showed features that would be expensive to include using conventional manufacturing methods.First Place: M.Arie Kurniawan IndonesiaThird Place: Sebastien Vavassori United KingdomFifth Place: Fidel Chirtes RomaniaSeventh Place: Andreas Anedda ItalyFigure 6.

Winning bracket designs1409Second Place: Thomas Johansson SwedenFourth Place: Nic Adams AustraliaSixth Place: Mandli Peter Hungary.Eighth Place: Piotr Mikulski PolandLessons LearnedAs already mentioned, many promising entries failed to fit within the original design envelope, which was required by the rules. This check could have been automated with immediate feedback to the entrants so that design modifications could have been made before the close of the contest.Both phases of judging were far more difficult than anticipated. Many of the light entries failed to meet the yield stress requirement, and it became clear that some of these entrants had not performed the analysis themselves. In the future, proof of a successful analysis should be a requirement for consideration.File size became a concern. Though most entries were of reasonable size, some were in excess of 300 MB.

In the future, proof of a successful analysis should be a requirement for consideration.File size became a concern. Though most entries were of reasonable size, some were in excess of 300 MB. Some of these large files were impossible to analyze in Phase I because they could not be meshed.Communication between GE and the community was difficult and improvements will be implemented in future challenges. GE’s evaluation became an unexpected bottleneck as judges were simply overwhelmed. An interesting suggestion evolved: Can the evaluation itself be somehow crowd-sourced, thus eliminating the need for judging in Phase I?ConclusionsThe Challenge created a high level of interest throughout the community and large number of innovative design concepts were proposed and exchanged. This was educational for all those involved.

This was educational for all those involved. A key question regarded the sustainability of such a crowd-sourced solution approach: Ten of the 700 entries in GE’s contest were awarded cash prizes, which left 690 entrants who were uncompensated for their efforts. Since many efforts were significant, would participation in similar future challenges result in similar worldwide interest? The answer is clearly “yes,” as interest in further contests is high.Several specific benefits to the Company were identified:Before the contest was announced, GE internal designers used proprietary design optimization tools to predict the outcome of the contest. Opening up the contest to the entire world community allowed GE to benchmark its own capabilities against hundreds of other approaches. It was enlightening to see the vast number of design innovations that could be applied to this rather simple part.

It was enlightening to see the vast number of design innovations that could be applied to this rather simple part. The GE team recognizes a considerable potential that may unlock creativity in more complex parts. One of the top ten contestants was hired as a GE employee under an internship program in Hungary, demonstrating to our Human Resources organizations that open innovation adds a valuable recruiting tool.In terms of dollars and fuel savings, implementation of the winning bracket will reduce the weight of a GE LEAP aircraft engine by 1.7 kg (3.76 lbm).

If similar weight reductions were to be applied to all flying aircraft worldwide, annual fuel savings from 12 to 22 million gallons of fuel (at a current value of $37M - $71M) could be realized.As a result of the success of the experiment, GE announced in May 2104 the activation of its global co-creation community FirstBuild.comvi – an online community dedicated to conceiving, engineering and building the next generation of major appliances for use in the modern home. Engineers, designers, scientists and home enthusiasts will participate in the development of breakthrough major appliances and solve engineering challenges. Participants will identify market needs, directly participate in the product development and watch via social media as ideas speed from mind to market at the FirstBuild™ micro-factoryvii in Louisville, KY.

GE opened the community to the public on May 15 with two challenges already in place.Developing a micro-kitchen: FirstBuild is participating in the NYC Economic Development Council's annual "BigApps" challenge that started May 7. FirstBuild will engage existing online communities as well as the FirstBuild.com community to help solve a growing problem in many metropolitan areas across the globe: increasing urbanization and population growth, which are driving the size of homes down and mortgages up. The challenge is to solve the design and engineering hurdles in creating a micro-kitchen that will maintain the style and functionality, but in a significantly smaller footprint than traditional kitchens. Indoor grilling: Grilling is a major part of the American lifestyle. A 2013 study shows that 80 percent of U.S. households own a grill or smoker. Yet, grilling can be dependent on the weather.

Indoor grilling: Grilling is a major part of the American lifestyle. A 2013 study shows that 80 percent of U.S. households own a grill or smoker. Yet, grilling can be dependent on the weather. FirstBuild is asking the global community of makers, designers and aspiring chefs to take the weather out of the equation by designing and engineering the ultimate indoor grill.In addition to these challenges, GE has more than 40 challenges under consideration forfuture engagement via crowdsourcing.Referencesi Brian Yanoff, et.al., Open Innovation in Precision Additive Manufacturing with Refractory Metals, RAPID 2014 Conference and Exposition, Detroit, MI, June 9-12, 2014. ii Wikipedia, the free encyclopedia. iii Southwest Airlines Co., 2012 Annual Report to Shareholders, p.43. iv http://en.wikipedia.org/wiki/Jet_fuel v http://grabcad.com/ vi https://firstbuild.com/ vii Microfactory, Inc. is a wholly owned subsidiary of GE, and FirstBuild is a registered trademark of Microfactory, Inc..

LCAST MOLDING WITH RAPID PROTOTYPED FUGITIVE MOLDSVikram R. Jamalabad, Philip J. Whalen AlliedSignal Research and Technology, Morristown, NJ 07962.John Pollinger AlliedSignal Ceramic Components, Torrance, CA 90509.Mukesh K. Agarwala, Stephen C. Danforth Rutgers University - Center for Ceramic Research, New Brunswick, NJ 08854.ABSTRACTA technique for the rapid manufacture ofceramic components has been developed using rapid prototyping to generate molds for the required components. The process entails the fabrication offugitive tooling using rapid prototyping techniques from which ceramic articles are formed using gelcasting. In the gelcasting process, the mold cavities are filled with a fluid suspension of ceramic powder which sets to a solid form through the polymerization of gelling additives and application of heat. The mold is carefully removed by dissolution or heat treatment leaving the intact gelcast part.

The mold is carefully removed by dissolution or heat treatment leaving the intact gelcast part. The "green" Computer aided gelcast part manufacturing of the tooling using solid freeform fabrication techniques allows for complex shapes to be manufactured with minimal tooling cost. The technique is ideal for the manufacture of ceramic parts in small batch conditions or for prototyping of Cost and time reduction of a magnitude can be functional parts in design cycles. achieved.is subsequently dried and sintered to full density.KEYWORDS: Solid Freeform Fabrication, Molds, Gelcasting, Ceramic Processing, Structural and Functional Ceramics.INTRODUCTIONSolid Freeform Fabrication (Rapid Prototyping / Manufacturing) is fast maturing into a means for the manufacture of structurally sound functional components. Until recently, only plastic and polymeric parts have been manufactured by SFF techniques.

Until recently, only plastic and polymeric parts have been manufactured by SFF techniques. 3D Systems' Stereolithographyl, the earliest of the techniques, is used for manufacturing plastic parts. Apart from design purposes, these parts are used as investment casting positives. Structural use of SLA parts is limited. Griffith et az2 have used ceramic suspensions in the SLA resin bath and have shown some progress in generating ceramic parts.Extensive work has been reported in the use of Selective Laser Sintering (SLS) by Bunnell et aP for the freeform fabrication of metal parts. Cima et at have demonstrated71the 3 Dimensional Printing (3DP) process for both, metal and ceramic parts.

Cima et at have demonstrated71the 3 Dimensional Printing (3DP) process for both, metal and ceramic parts. Fused Deposition of Ceramics (FDC), based on Fused Deposition Modeling (FDMTM) has been shown to be feasible (Agarwala et aZ5 In addition, some novel technologies that are in early stages of development (SALD, by Zong6 and Marcus) could lead to the manufacture of structural components.).Indirect methods of manufacturing structural components using solid freeform fabrication have been used since the inception of these techniques. Part positives for investment casting can be made by almost all SFF processes. Investment casting wax parts (Stratasys Inc's FDMTM machines), paper parts (Helisys Inc's LOMTM machines), polymeric parts (Sanders-Prototype Inc's MM-6PRO Model Maker™, MIT's 3DP machines) and photopolymeric parts (3D Systems Inc's SLATM series machines) have been utilized.

Ceramic slurry (usually a refractory like silica) is set around the part positives before the parts are melted away to leave a cavity for casting.The development of complex shaped ceramic structural components and devices is severely hampered by the high cost of ceramic prototype manufacturing and the overheads of short production runs. Machining of sintered, fully dense ceramic articles is extremely hard, time consuming and expensive. This has led to the development of many net shape fabrication techniques. Injection molding and slip casting are two of these techniques. Both require some form of precision tooling, effectively driving up cost.The overall cost of a component will, in most cases, be driven by the cost of the tool fabrication. Time consuming and expensive machining operations' are required for The necessity of this tooling also precludes the use of iterative tool preparation.

Time consuming and expensive machining operations' are required for The necessity of this tooling also precludes the use of iterative tool preparation. approaches in design, since changes in component design require significantly high levels of re-investment in time and capital. For injection molding, high quality water-cooled dies are required. At the same time, these need to be rapidly dis-assemble to eject the part. Very large cross-section parts also cannot be manufactured due to binder burnout problems. Complex porous tooling is required for slip casting. Variations in particle distribution lead to warpage and an inability to hold tolerances.RAPID PROTOTYPED MOLDING FOR GELCASTINGCeramic gelcasting is a process whereby a ceramic slurry containing appropriate monomers is poured into a tool and then "gelled" through a catalyzed reaction. This gelation polymerizes the monomer resulting in a relatively rigid solid.

This gelation polymerizes the monomer resulting in a relatively rigid solid. Metal and ceramic particles in the suspension can be gelled in this manner. Low cost multi-part metal (aluminum) tooling is used for ceramic gelcasting. While this works for relatively simply shaped parts, for complex parts, fugitive tools or molds which are removed by dissolution can be used. Gelcast tooling must have several key features. The tool must be stable at the temperatures required for gelling (40 to 60 C). the tool must be structurally sound to support the ceramic suspension during the gelling process.In addition,72---Conventional Fugitive ToolingPrepare drawings or CAD of Metal tooling....CAD file of partRP Fugitive Tooling• of tooling,Prepare .SLC fileMachine metal tooling ..... Cast soluble wax positives using toolingFabricate pattern fixturingGenerate .SML file and check for integrity..Assembl::'ax sections.

Cast soluble wax positives using toolingFabricate pattern fixturingGenerate .SML file and check for integrity..Assembl::'ax sections . / into complete positivePrepare non- water soluble wax tooling: - Dip soluble wax positive into non-soluble wax - Dissolve soluble wax leaving finished toolingFabricate wax or ABS tooling by FDMMultipart Tool: Glue sections together (if required)Figure 1. Process for fabrication of Fugitive ToolingThe fabrication of conventional fugitive tooling is a very labor intensive process with long lead times. Figure 1 shows how the fabrication of fugitive tooling can be significantly simplified through the use of solid freeform fabrication techniques. Using a the cost large ceramic turbine nozzle doublet component (Figure 3) as an example, savings of using rapid prototyped fugitive tooling was examined (Tables 1 & 2).

Using a the cost large ceramic turbine nozzle doublet component (Figure 3) as an example, savings of using rapid prototyped fugitive tooling was examined (Tables 1 & 2). The results show that significant time and cost savings can be realized for both the production of the i~itial part and subsequent components.Mold design is a relatively mature area of work. Rapid prototyping procedures provide an efficient means of manufacturing molds directly from the envisioned CAD models. Minimal tooling investment is required, apart from the prototyping machinery itself. Complex shapes with non-extractable surfaces and blind holes can be designed with ease. The direct fabrication of fugitive tooling from a CAD file significantly reduces the time and cost of fabricating complex net shape ceramic components by eliminating excessive machining. The additional benefit is that changes in the component design can be easily accommodated without additional capital equipment investments.

The additional benefit is that changes in the component design can be easily accommodated without additional capital equipment investments. The details of fabricating gelcast tooling using rapid manufacturing techniques and using that tooling for ceramic gelcasting, shown schematically in Figure 2, is described below.1. The CAD model of the component is scaled up to allow for shrinkage. 2. A mold is designed for the component with the following features -Part orientation for minimal mold overhangs and elimination of air pockets.73A solid lining for the mold cavity, created by conformally offsetting the component surface and building a part "perimeter."A honeycombed mold support structure behind the mold cavity. The honeycomb structure increases the ratio of surface area to volume of the mold material and results in more efficient dissolution.3.Add appropriate reservoirs, gatings and risers for easy filling of the slurry.

If a mold design has large overhangs or sections in which support structures are hard to remove, multi-part molds are built and assembled using appropriate adhesives. Locating holes are built into the mold to enable accurate alignment of the parts.4. Support structures are added as required. These are preferably of materials dissimilar to the part material.5. The molds are built on the appropriate RP platform of acetonitrile butadiene styrene (ABS) or investment casting waxes.6. Molds are tested for integrity by filling with water and holding for a period of time. 7. The ceramic gelcasting slurry is poured into the cavity of the mold so that all air pockets are removed.8. Gelation is then achieved by elevating the temperature. 9. The entire assembly is immersed in a bath of solvent (acetone, alcohol, toluene) and In the case of wax, elevated temperatures can be used to melt the mold is dissolved. away the mold.10.

9. The entire assembly is immersed in a bath of solvent (acetone, alcohol, toluene) and In the case of wax, elevated temperatures can be used to melt the mold is dissolved. away the mold.10. The formed ceramic component that is left behind is dried and sintered.(a) CAD(b) Processing(c) Mold(d) Final PartFigure 2: Schematic of the Gelcasting Process for a Ceramic Turbine Blade.The cost advantages of this technique are manifest in the following: (a) Minimal machining is required in the fabrication of a complex ceramic component. (b) Rapid generation of molds with varying component parameters enables quick turnaround in iterative part design. (c) Defects and part rejection caused by problems with mold disassembly are minimized.

(c) Defects and part rejection caused by problems with mold disassembly are minimized. (d) Doing away with physical means of removal ensures that thin sections and delicate protrusions can be manufactured.EXAMPLETo demonstrate the feasibility of using rapid prototyped fugitive tooling to fabricate complex ceramic components the entire process was exercised on a ceramic turbine74the Office of Naval Research.75benefits are realized by time savings by eliminating costly tool machining procedures. The technique enables the use of iterative design and the fabrication of parts with unique structures (complex protrusions, thin sections).ACKNOWLEDGMENTSThis work was sponsored by the Defense Advanced Research Projects Agency and the Office of Naval Research under contract # N00014-94-0115. Thanks are due to Barry Draskovich and the gelcasting team at AlliedSignal Ceramic Components, the engineering group at Stratasys Inc., Ed Pankowski at FabTek Inc.

Subsequent iterations of the same mold would take 71 to 76 hours at a cost of $ 990 to $ 1265/mold.Table 2. Cost Analysis for the Fabrication of Standard Wax Tooling for Gelcasting the Ceramic Turbine Nozzle Doublet.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperGENERAL RULES FOR PRE-PROCESS PLANNING IN POWDER BED FUSION SYSTEM – A REVIEWTan Pan1, Sreekar Karnati1, Frank Liou11Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractPowder bed fusion (PBF) is one of the current additive manufacturing techniques that can fabricate almost fully dense functional metal components. Through a layer by layer fabrication methodology, complex geometries to meet the requirements of aerospace, automotive, biomedicine industries, etc. can be produced. The success of a build largely depends on having a flawless pre-process planning, including build orientation selection, support structure optimization, process parameter chosen, etc., which closely relates to the quality of the final products.

Geometric inaccuracy and poor surface quality can occur due to a bad build plan. This review presents the crucial general planning rules for the build process. Build orientation selection, support structure optimization, and process parameter chosen in terms of residual stress reduction are the mainly concerns, which have been surveyed and discussed. The overall objective of this work is to help setup build plans that can ensure precise dimensions and high surface quality among the built components.IntroductionPowder bed fusion (PBF) employs a manufacturing process by sintering or melting the raw material with a heat source (usually a laser or electron beam) then cooling and consolidating the material layer by layer to allow the fabrication of light-weighted components. The layered manufacturing methods enable the fabrication of the small-size components with the complex geometry.

The layered manufacturing methods enable the fabrication of the small-size components with the complex geometry. The geometry complexity is of difficulty for the conventional manufacturing techniques to realize.However, some consequences of an unsuccessful pre-processing plan hold back the prevalence of PBF system. Given that the fabricated products usually possess a poor surface finish [1,43-47], which needs machining during post-process. Besides, the strong support structure is necessary to be applied to constrain the overhang or inclined angle structure so as to improve the stiffness of the supported structure and dissipate the redundant heat efficiently, otherwise most time, the build will fail [13,14]. Nevertheless, the conventional support is usually a volume column underneath the supported material.

Nevertheless, the conventional support is usually a volume column underneath the supported material. The additional build of the support structure leads to the extra cost of material and time to design during pre-processing, manufacture during the build processing and remove at post-processing [16]. The residual stress because of improper use of process parameter gives rise to the deformation, distortion, and delamination or even build failure [30]. The origin of the residual stress in the layered manufacturing process is temperature gradient mechanism. That the large thermal gradient comes from the rapid heating and relatively slow cooling induces stress inside the material. The stress is not allowed to fully release by an elastic1161(stress less than the yield stress of the material) or plastic deformation (stress more than the yield stress of the material) due to the fast solidification, so part of the stress remain inside the components.

[29-30,32,38]Proper pre-process planning is one of the inspirations that could liberate the build process from those limitations. In this review, a pre-process planning in terms of the optimization of build orientation, the selection of the support structure as well as the reduction of residual stress from changing the process parameter are presented as a guideline to achieve a better geometric accuracy, surface finish and save more build time and material.Pre-Process Planning SetupBuild orientation selectionThe feature and objective of the build need to be specified first so as to obtain an optimized build orientation [24]. Shape features need to be identified to determine the manufacturability of the Additive Manufacturing (AM) processes, which related to the constraints of the layer-wise manufacturing process and the contact between the components and the machine. The shape features are machine and build geometry dependent.

The shape features are machine and build geometry dependent. For instance, for different manufacturing processes based on powder bed fusion (PBF), such as Selective Laser Melting (SLM), Selective Laser Sintering (SLS), and Electron Beam Melting (EBM), there are some special rules regarding the contact between the fabricated parts and the wiper or recoating blade to prevent the extreme rise of the edge of the part [25].a. No long parallel contact to reduce the force induced by a wiper or recoating blade to a single part. It can be solved by rotating a small angle of the part (Figure 1). For multiple parts in one build, the parts could not be positioned along the wiper to eliminate simultaneous initial contacts [25].b. The inclined overhang or support structure need to face away from the wiper or recoating blade so as not to exacerbate the upward trend of the part deflection on account of the force induced by the movement of the wiper [25].c. Avoid too large aspect ratio.

Avoid too large aspect ratio. The force from the wiper will bend the thin, tall part eventually[25].Figure 1: Build orientation chances by a small angle rotation1162General objectives that need to be defined in terms of the designer preference and requirement to achieve the better functional products. The main general objectives include surface quality [24,26,35], part accuracy [24], support volume [24,26], mechanical properties [24], build time (build height) [24,26], build cost [24], post-processing [24], favorableness of AM feature [24], and other potential attributes [26]. Given that the surface finishes of the upward and downward facing surface of a fabricated overhang are distinct [48]. The build orientation of the overhang needs to be changed upside down if the user requires a better surface finish of the down skin, as shown in Figure 2.After the features and objectives are specified, two main tasks need to be figured out [21,24].

First, select a set of the alternative orientations from the infinite build orientation for a part based on the shape feature. Rule or knowledge-based method, sampling or listing method could be employed, listed in Table 1. Genetic algorithm and Response Surface Methodology are the two representatives in sampling or listing method. Then refine the alternative orientation to get rid of duplicate orientations. The second step is to determine the optimized orientation in terms of the identified general objectives or attributes applying multi-criteria decision-making method [20,23,24].Figure 2: Build orientation changing with a key featureZhang et al. [21] use a feature and rule-based method to generate finite orientation sets for every single part. Then apply the multi-criteria decision-making method to rank and weight the criteria to determine the optimized orientation for multi-part.

Then apply the multi-criteria decision-making method to rank and weight the criteria to determine the optimized orientation for multi-part. Byun and Lee [19] determine their optimized build orientation from all the possible candidates by genetic algorithm (GA). They rank and weight the importance of each criterion (surface roughness and build time as a priority in this case) to identify the optimized orientation. Zhang et al. [20] point out a two-step solution to determine the optimized orientation for multi-part production. The single part orientation sets can first be selected regarding the key feature of the user preference. Then the optimized orientation for multi-part can be identified by setting the global objectives and applying the GA. The authors of [28] propose a software using Matlab with an optimization algorithm to determine the build orientation with least support volume, which could be applied to both single part and multipart.Ahsan et al.

[26] further develop an analysis for concurrent determination of both optimal build orientation and tool-path direction. The analysis combines the basis of geometry, the part attributes, and genetic algorithm. Recently, Langelaar [27] provides a method, which could be applied to optimize the part geometry, build orientation and support layout simultaneously.1163Table 1: The method used to optimize the orientationSupport structure for overhang problemThe supports used in building process are significant structures designed as sacrificial structures so as to reinforce the resulting components. The designed support structures are well needed under the following situation in SLM process [48]:I.The overhang plane (length more than 1mm) with an angle of less than 45 degrees with regard to (w.r.t.) the horizontal plane. Otherwise, either a relatively low density or a rough finish of the down skin surface could happen.

Otherwise, either a relatively low density or a rough finish of the down skin surface could happen. The length of the plane within 0.3mm to 1mm can self-support without support structures [48].II.Local minima refer to any areas of the component that are not connected to the layer below, which will need the support to anchor them. If not, the first layer of the local minima is possible to be displaced by the wiper, which may result in inaccurate geometry, or even failure [48].III.Lateral holes or arches on the side surface of the component with the radii more than 5mm will also need the support in order to prevent geometry distortion from happening [48].Support structure optimization is of importance. With a proper selection of the support structure, material and build time could be saved, better surface finish could be achieved, and the support could be easier to be removed. The reviewed support optimization strategies are mainly categorized into two types.

The reviewed support optimization strategies are mainly categorized into two types. One is designed in terms of the support geometry selection. The other is optimized by changing the process parameter for both downward facing surface and support.1164These two support design strategies are listed in Table 2.Figure 3: Conventional solid column support after EDMI.Support GeometryTo achieve simple support and fast computation, the authors of [13] present a model and algorithm to optimize the shape and topology of the support. The shapes of conventional support structures are usually solid column like Figure 3, which are waste of the material and build time. So instead of a conventional solid column, lattice support structure is a better choice. Hussein et al. [42] study the two types of advanced lattice support structures, gyroid and diamond.

So instead of a conventional solid column, lattice support structure is a better choice. Hussein et al. [42] study the two types of advanced lattice support structures, gyroid and diamond. They point out that in the best case, only 8% relative volume is used with gyroid lattice support structure by changing the size and struts of the unit cell, which means 92% of the powder can be saved and reused. The diameters of the struts of the support structure and the distances between the adjacent points connected to the down skin surface are two determinant factors to influence the manufacturability. The authors of [12] also present the feasibility of the gyroid and diamond cellular support structure with different size of unit cell and volume fraction. The manufacturability and performance of the gyroid type cellular support structure made by SLM are evaluated by Yan et al. [6].

The manufacturability and performance of the gyroid type cellular support structure made by SLM are evaluated by Yan et al. [6]. The conclusion shows unit cell size between 2 to 8mm can be fabricated properly without additional support and the geometric accuracy of the produced structure is sound.The authors of [11] present a pure mathematical 3D implicit functions for design and optimize the cellular support structure. Vaidya and Anand [17] propose an approach to generate a cellular support structure. They combine the space filling hollow cellular support with Dijkstra’s shortest path algorithm to create the support with minimal volume fraction.Another commonly used geometry of the support structure is teeth support [7]. Poyraz et al. report a design of the teeth support then conclude that the tooth parameter is less important than the hatching distance which greatly determines the extent of the distortion of the overhang.

report a design of the teeth support then conclude that the tooth parameter is less important than the hatching distance which greatly determines the extent of the distortion of the overhang. Calignano [8] investigates the optimized teeth parameter for AlSi10Mg and Ti6Al4V in terms of process efficiency and effectiveness by selective laser melting and demonstrates the importance of build orientation selection to reduce the amount of support and the possibility of distortion. Figure 4 illustrates the teeth support below the fabricated part. The authors of [7] propose the easiness of the removal of the teeth support during post-process due to the weak tip of the teeth.1165Figure 4: Teeth support geometryIn addition to the lattice and teeth support structure, there are still some effective support structures with different geometry developed.

Gan and Wong [2] investigate three support structures, inverted ‘Y’ shape, ‘Y’ shape, and pin and came to the conclusion that the relatively levelled thin overhang plates can be achieved with the help of the inverted ‘Y’ supports with only 2.2% overhang-support contact area. The authors of [10] come up with a point contact support which can dissipate heat efficiently with the minimal contact area, which is also easy to remove during post-processing.Recently, a new concept of the support structure realizes in Electron Beam Melting (EBM) process, which is developed by Leprince-ringuet and Vignat [3]. They point out a contact-free support which could not only be removed without any mechanical tool but reduce the extent of distortion efficiently. But the problem is that the contact-free solid column which could not be recycled is still a waste of material and production time.

But the problem is that the contact-free solid column which could not be recycled is still a waste of material and production time. The optimization is proposed to add a constraint at the front of the part end, which can also effectively prevent the overhang from curling at the edge meanwhile maintain the ease of support removal and save the material significantly. Cooper et al. [4] also come out with the idea of contact-free support as a heat sink. They demonstrate the applicability of the heat support with no or little post-processing by both simulation and experiment and figure out that the effectiveness of the support is sensitive to the gap distance between the downward facing skin and the support and thickness of the support. Li et al.

Li et al. [16] also propose a lightweight and support-free design method to design a support.II.Process parameterBesides the optimization of geometry of the support structure, altering the process parameter of the downward facing surface and the support also helps. Mertens et al. [9] propose an optimized laser power and scan spacing for downfaces of the horizontal and inclined overhang to achieve better performance without support. However, this design is just for short structure while the extra support has to be provided for long downfacing structures. Cloots et al. [10] demonstrate that the support structure can be eliminated for the overhang with an angle of more than 20°w.r.t horizontal plane. They change the hatching space, scan angle and scan speed to reduce the heat input and achieve the build without support while still retaining acceptable density and surface roughness. Jhabvala et al.

Jhabvala et al. [15] present that support structure could be built using pulsed laser accounting for easy removal and time-saving. Bobbio et al. [18] use a relatively lower energy input to build the support structure than the solid part to achieve a low structural strength, which could be of benefit to ease the support removal during post-process.1166Table 2: The strategies of different support geometries and process parameterProcess parameter chosen in terms of residual stress reductionThe consequence of the high residual stress may result in a severe build failure like delamination in Figure 5. Therefore, it is necessary to evaluate and control the magnitude of residual stress during the build process to prevent the curing or warping happening. The authors of [32] give a detailed review of residual stress and further a future study direction.

The authors of [32] give a detailed review of residual stress and further a future study direction. Many researchers report the methods they use to measure the residual stress and analyze the effects of different factors on the residual stress.Figure 5: The consequence of large residual stressKruth et al. [31] provide a method called bridge curvature method used to assess the angle of curvature α of the bridge structure made by SLM to imply the magnitude of the residual stress, shown in Figure 6. Thus, this method can be employed to assess and compare the extent of the influencing factor on residual stress. Zaeh and Branner [30] analyze the residual stress by using neutron diffractometer. Safronov et al. [36] put forward that the top surface of the part could be concave on account of residual stress induced shrinkage. Thus, the magnitude of the residual stress1167is related to the curvature radius R.

Thus, the magnitude of the residual stress1167is related to the curvature radius R. The authors of [37] evaluate the residual stress by hole-drilling strain gauge method and conclude the residual stresses in some area of the part exceed the yield stress of the material, which indicates the distortions in the specimen result from the residual stress. Liu et al. [38] measure the residual stress by X-ray diffraction and propose the stress perpendicular to the scanning direction are much smaller than that parallel to the scanning direction. In addition, the peak values of residual stress always exist at the beginning of scanning tracks, which means the deformation or distortion are easier to happen at the edge of the part. The authors of [39] also demonstrate that the greatest stress occurs parallel to the scan vectors since there is larger thermal gradient parallel to scan vector using a thermo-mechanical model.Figure 6: (a) Before, (b) after removal from the base plate.

The geometry of the test partsMercelis and Kruth [29] propose the distribution of the residual stress, which contains large tensile stresses at both top and bottom of the final part while intermediate compressive stress in between. They also investigate the parameter influencing the magnitude of the residual stress and the method to reduce the residual stress; details are generalized in Table 3. The application of island scan strategy generates lower residual stress than that of longitudinal scan direction or alternating scan direction. Zaeh and Branner [30] also report the parameters which effect on residual stress and deformation. They conclude that the residual stress could be a function of the scanning direction, cantilever thickness, layer thickness and initial platform temperature. Further, Ali et al.

They conclude that the residual stress could be a function of the scanning direction, cantilever thickness, layer thickness and initial platform temperature. Further, Ali et al. demonstrate that pre-heating the bed to 570°C for Ti6Al4V could significantly reduce the residual stress and improve the yield strength and ductility of the final components [34]. The authors [39] indicate the longitudinal stress increases with scan vector length and a non-uniform anisotropic stress field is generated during the layer by layer manufacturing process. Vasinonta et al. [40] employ a thermo-mechanical model with a thin-walled structure to quantify the effect of wall height, scan speed, laser power and the preheating of the substrate on the size of the melt pool and the residual stress. Two process maps are applied to predict the melt pool size and maximum residual stress.

Two process maps are applied to predict the melt pool size and maximum residual stress. The process parameter can be determined using this method to optimize the melt pool size while controlling the residual stress.1168Table 3: The effect of different factors on residual stressConclusionTo keep the build from distortion and deformation, and ensure the quality of the final components, a proper build setup need planning in powder bed process. In this review, the setup rules with regard to orientation selection, support structure optimization and residual stress reduction are concluded as follows:1. Build orientationThe rules of orientation optimization consider the geometry feature and wiper interaction. The orientation needs to be selected in terms of preventing the extreme rise of the edge of the part on account of the interaction between the build part and the wiper movement during the build process.

The alternative orientation sets can be generated by either rule (knowledge) based method or sampling (listing) based method. After general objectives of the build are specified by the user, the optimal orientation could be selected from the sets.2. Support structure optimizationThe structures with the angle of a plane less than 45°w.r.t the horizontal plane, with local minima or with lateral holes with large diameter need to be supported. To save the build time and minimize the support material, different geometry of the support structure which can still self-1169support could be applied. In addition, the support structure could be built using pulsed laser or low energy input to ensure the ease of the final support removal. The process parameter with lower energy input could also be employed to build the downward facing surface to achieve a better surface finish without support.3.

The process parameter with lower energy input could also be employed to build the downward facing surface to achieve a better surface finish without support.3. Process parameter chosen in terms of residual stress reductionThe effect of different factors, including the number of layers, the thickness of the baseplate, the scan pattern, and energy input, etc., on residual stress, are mainly listed in Table 3.AcknowledgmentsThe support provided by Intelligent Systems Center (ISC) and Center for Aerospace Manufacturing Technology (CAMT) at the Missouri University of Science and Technology is appreciated.Reference[1] Calignano, F., Manfredi, D., Ambrosio, E. P., Iuliano, L., & Fino, P. (2013). Influence of process parameters on surface roughness of aluminum parts produced by DMLS. The International Journal of Advanced Manufacturing Technology, 67(9-12), 2743-2751. [2] Gan, M. X., & Wong, C. H. (2016). Practical support structures for selective laser melting.

nerating Topological Informa*ion from a "Bucket of Facets" Stephen J. Rock Michael J. WoznyRensselaer Design Research Center Rensselaer Polytechnic Institute Troy, New York 12180AbstractThe STL de facto data exchange standard for Solid Freefonn F*brication represents CAD models as a collection of unordered triangular planar facels. No topological connectivity infonnation is provided; hence the term "bucket of facet~." Such topological information can, however, be quite useful for performing model v~lidity checking and lfhis paper discusses speeding subsequent processing operations such as model slicing. model topology and how to derive it given a collection of unordered tri,ngular facets which represent a valid model.1 IntroductionComputer Aided Design (CAD) model data is frequently pas~ed to various Solid Freeform Fabrication processes using the STL polygonal facet repre~entation [1].

Facet models represent solid objects by spatial boundaries which are defin~d by a set of planar faces. This is a special case of the more general Boundary Represent,tion which does not require object boundaries be planar [2]. In general, the term facet isl used to denote any constrained polygonal planar region being used to define a model bottndary; however, in the Solid Freeform Fabrication (SFF) community the term facet is typlically understood to mean triangular facet. Representing models using triangular facets ha~ both good and bad points [3]. Facets do provide a "greatest common denominator" geo~etrical form for data exchange between many CAD systems and SFF processes. Non-C~D scalar field data, such as that from CT imaging, can also be used to generate facet mqdels [4].

Non-C~D scalar field data, such as that from CT imaging, can also be used to generate facet mqdels [4]. However, facet models are generally only an approximation of mathematically prepise CAD models.Precise CAD models must be tessellated, where defining P'lodel surfaces are subdivided into planar facets, to create polygonal facet models [5]. lAs model precision demands become more stringent, the number of facets required to adeq*ately approximate a model will increase. Model tessellation should yield a set of facets wpich define a closed region representing the material boundary of a part. Unfortunately, m~y commercial CAD system model tessellators are not robust, and sets of facets which qo not define closed regions result. This missing facet problem is particularly prevalent wh~re surfaces intersect in the original CAD models. A set of facets which, when assembled,lforms a solid object with holes in its surface is incomplete and is termed an invalid mqdel.

A set of facets which, when assembled,lforms a solid object with holes in its surface is incomplete and is termed an invalid mqdel. In addition to missing facets, other causes of model invalidity exist. They include errqrs due to numerical round-off, missing data, altered data, and sometimes the presencle of extraneous or redundant data.The de facto industry standard STL model representation defin~s models as a set of triangular facets [1]. Unfortunately, these facets are stored independeptly, as if each facet were created and tossed into a bucket with no particular ordering and }vithout information relating a given facet to any other facets in the bucket. Since many ~AD systems fail to generate valid facet model tessellations, it is necessary to perform moqel validity checking before subsequent processing operations are undertaken. Given onlyl the data in an STL file, performing model validity checking is computationally expensiive.

Given onlyl the data in an STL file, performing model validity checking is computationally expensiive. Attempting to© 1992 RPI Rensselaer Design Research Center. All Rights Reserved.251determine the relationships, or topology, between model facets from the "bucket of facets" is the first step in performing validity checking. The resulting topological information is important for use in subsequent processing operations such as model repair, model slicing, and finally during the scan conversion operation.2 TopologyTopology describes the connectivity relationships between various geometric entities [2]. A facet can reference the three edges which bound it. Each edge can reference the two vertices which define it. Topological connectivity relationships are not limited to individual facets. For instance, a facet can reference the three facets which share edges with it. An edge can reference not only the two vertices which define it but also the two facets which share it.

For instance, a facet can reference the three facets which share edges with it. An edge can reference not only the two vertices which define it but also the two facets which share it. Vertex points can contain connectivity information to all edges or faces which share it. Such references are all examples of topological connectivity information.It is important to consider two topological classes of boundary representation models: manifold (two-manifold) and non-manifold. A two-manifold is defined as a two dimensional, connected surface where each point on the surface has a neighborhood topologically equivalent to an open disk [6]. In a two-manifold, every edge in the model is shared by two and only two facets. This is the case for most facet models where only the facets representing a part's spatial boundaries exist. One side of a facet is directed toward part material, and the other is directed away from it.

One side of a facet is directed toward part material, and the other is directed away from it. The spatial boundaries of a facet model are expected to have a distinct "inside" and "outside" which is consistent across all facets defining the model boundary; such a model is termed orientable [2]. Non-manifold conditions occur where, for example, two distinct enclosing volumes share one facet or a set of facets as a common boundary [7]. In this case, the shared facet no longer has a clear "inside" or "outside". Both sides of the facet are surrounded by part material. This typically occurs when facets representing multiple solids which are tangent along some boundary are not properly delimited as belonging to individual solids in an STL fIle.3 Benefits of Topological InformationThe STL format represents facet models with nearly the minimal information necessary to define a solid object.

Each facet, along with its normal, is specified explicitly and no topological connectivity information is provided. This "bucket of facets" approach to model representation has many limitations, both with respect to ensuring valid models and subsequent processing. 3.1 Validity CheckingWhen no topological information is provided, model validity checking involves computationally expensive searching operations. If model topology were available, validity checking would be a much simpler and efficient operation. 3.2 Model RepairWhen invalid models are encountered, topological information is useful for attempting to repair the models. Such information makes it readily apparent when greater than two facets share a single model edge. In the case of model holes, it is important to know how the facets surrounding a hole are connected and this too can be easily determined given topological information.

In the case of model holes, it is important to know how the facets surrounding a hole are connected and this too can be easily determined given topological information. 3.3 Subsequent ProcessingFacet model slicing performance can also benefit from topological information which makes it possible to march from facet to neighboring facet performing simple edge/plane intersection calculations [8]. This same topological information can be passed to subsequent processing phases, such as scan conversion, which occur after slicing.2523.4 Model RepresentationFinally, model topology generation capability has facilitated the development of a richer facet model representation format [9]. Storing topological information with a facet model, although increasing the information content of the model, reduces the volume of data required when compared to an equivalent model represented in STL format.

The net result is a more robust representation with less data.4 Topology Reconstruction ConceptsReconstructing model topology given a "bucket of facets" is basically a searching operation. Entity relationships must be found by searching the unordered model data. These relationships, or topology, must then be stored for later use. The conceptual steps for producing such topology will now be discussed; however, implementation details, such as the searching algorithms or structures used, will be dealt with separately after the topology reconstruction concepts are understood. 4.1 Vertex MergingEach facet is defined by three vertices whose coordinates are explicitly specified. The first operation performed when reconstructing model topology is vertex merging. Here, equivalent, explicitly specified vertices .are replaced bya single entry in a list of unique vertices. Each face then references three vertices in the list instead of being defined by actual vertex values.

By using topological information, a unique definition of each vertex can be referenced by the facets using pointers. If a pointer consumes four bytes, this will reduce the memory required to represent the four facets shown to 168 bytes (5 vertices x 3· doubles/vertex x 8 bytes/double + 4 facets x 3 pointers/facet x 4 bytes/pointer). This figure shows a significant storage savings where only four faces meet. However, in a real model which is closed and likely more complex, many facets will share each single vertex. This will result in even greater savings by using topology and representing each vertex uniquely.253The final benefit of vertex merging is that vertices within a predetermined numerical round-off tolerance of each other can be easily merged, and this can be used to overcome errors introduced by inconsistent numerical round-off. Such errors can occur when a slightly different sequence of mathematical operations is used to calculate the same vertex value.

Such errors can occur when a slightly different sequence of mathematical operations is used to calculate the same vertex value. For example, given three finite precision binary numbers A, B, and C, the operation A + B + C may not produce the same result as the operation A + C + B due to rounding or chopping errors [10]. Figure 4.2 provides a two-dimensional example of how vertex merging can also remove facets smaller than the size of the numerical round-off tolerance.LI-_ II I I IL_------------- ----~ C Numerical RoundoffEquivalence RegionsFigure 4.2 - Facet Removal by Vertex MergingThe facet defined by the three vertices shown is smaller than the specified numerical round-off tolerance. The squares drawn with dashed lines and centered on each of these vertices indicate the equivalence regions defined by the numerical round-off tolerance setting. Since the vertices are within the equivalence regions, they are considered equivalent and should be merged into one vertex.

Since the vertices are within the equivalence regions, they are considered equivalent and should be merged into one vertex. This is illustrated by the three arrows pointing toward the center of the left hand figure. Notice that as the three facet edges collapse, the remaining edges of the three adjacent facets will approach each other.The figure on the right shows what happens to the model after the three equivalent vertices have been replaced, by merging, with one new vertex. The three facets which were adjacent to the facet removed by vertex merging now have two identical vertices (the vertex resulting from the merging operation). Note that this creates a degenerate condition where three of the facets surrounding the merged facet have effectively collapsed to an edge. The facet topology is still present; however, two facet vertices are the same in each of the degenerate facets. 4.2 Face & Edge CreationA face entity must be created to represent each facet in the model.

4.2 Face & Edge CreationA face entity must be created to represent each facet in the model. Since vertices are represented uniquely, each face maintains references to the three vertices which define it. Each face entity also carries with it a facet normal. This aids in determining which side of the face is inside the model; however, this information could also be computed if it were not given, provided the right-hand rule for vertex definition is adhered to [1].254Similarly, it is possible to create edge entities to represent model edges. Each edge can reference the two vertices which define it and the two faces which share it. Edge information is not provided by STL format files, but it can have utility when slicing facet models and can be represented using alternative rue formats [3]. The following section will show that it is appropriate to wait until face relationships have been determined before creating model edges.

The following section will show that it is appropriate to wait until face relationships have been determined before creating model edges. 4.3 Determining Face & Edge RelationshipsWhen face entities are created, they not only contain references to the defining vertices and normal, but they also contain unassigned references to the three adjacent faces and corresponding edges. Figure 4.3 provides a graphical depiction of adjacent face and edge references .Figure 4.3 - Adjacent Face and Edge References These references will be assigned as model topology is determined. After all face relationships are determined, the relationships between model edges can be determined using the face relationship information.Face relationships are determined by searching. For each face in the model, searching is performed to determine what other faces share two common vertices. The existence of a pair of such faces defines an edge.

For each face in the model, searching is performed to determine what other faces share two common vertices. The existence of a pair of such faces defines an edge. When such a face is located, the adjacent face references of each face corresponding to the shared edge are cross-referenced. This establishes one topological relationship between two faces. This process is repeated until all adjacent face references are set. Notice that this assumes each model edge is shared by exactly two faces; invalid models exist where this is not the case.After defining all face adjacency relationships, it is possible to define model edges. The three vertices which define each face and the. three adjacent faces of each facet are known. Edge topology can be derived by using this information. The three edges of each face, defined by the face's vertex pairs, can be added to an edge list if they do not already exist.

Edge topology can be derived by using this information. The three edges of each face, defined by the face's vertex pairs, can be added to an edge list if they do not already exist. When an edge is added, the vertices which define it are known, so its vertex references can be set accordingly. One of the faces which shares the edge is immediately known, as this is the face being used to reference the edge's defining vertices. The. other face which shares the edge is also known after the face adjacency relationships have been determined. By •• sequencing through all faces in the model after the face adjacency relationships are known, complete edge topology can be created without additional searching.2555 Data Structures & AlgorithmsThe fundamental operations used to reconstruct model topology have been discussed, but few operational details were provided.

If brute force approaches to vertex merging and adjacent facet searching are used, computational cost will quickly become prohibitive as model complexity increases. Consequently, care must be taken to ensure the data structures and algorithms employed facilitate efficient searching.Facets are read sequentially from a data file with no particular ordering. Each time data representing a facet is read, a face entity is created. Three vertex entities, each of which may already exist or need to be created, are referenced by each face entity. A linear list, although a poor structure for sorting and searching, can be used to store the face entities because they are not searched directly. Vertex entities must, before they are created, be tested to ensure an equivalent vertex does not already exist. If it does, this equivalent vertex should be referenced by the newly instantiated face entity instead of creating a new vertex entity and referencing it.

If it does, this equivalent vertex should be referenced by the newly instantiated face entity instead of creating a new vertex entity and referencing it. A searching operation is required to perform this equivalence testing, and it is repeated for each model vertex. Repeatedly searching, and maintaining in sorted order, a linear list of vertex entities is a costly proposition. An alternative to the linear list data structure should be used. 5.1 The AVL Tree for Vertex MergingWith the goal of vertex merging in mind, a linear list is clearly an unacceptable solution for storing the unique vertices. Each time a vertex entity is added, the complete list would have to be sorted and/or searched. Recall that one requirement for vertex merging is that all equivalent vertices be merged. This suggests that for a vertex entity to be added, there must be no existing equivalent vertices. Consequently, a range of vertex values in the neighborhood of the search key must be tested.

This suggests that for a vertex entity to be added, there must be no existing equivalent vertices. Consequently, a range of vertex values in the neighborhood of the search key must be tested. This makes a hash table an unattractive storage structure. A sorted array could be used, but because vertex data is continually being added as each model facet is read, repeated insertions into the array and the block moves which result make this an unattractive solution. A binary search tree could be used; however, the input data ordering is unknown. In the degenerate case, this could be equivalent in performance to searching a linear list of vertex values, and this is unacceptable.A balanced binary search tree, termed an AVL search tree, overcomes these limitations [11]. It never degenerates to the equivalent performance of searching a linear list because its balance is maintained as each element is inserted in it.

It never degenerates to the equivalent performance of searching a linear list because its balance is maintained as each element is inserted in it. It also meets the requirements for being able to traverse through the data and efficiently allow range searching, which is used to determine equivalence within a set numerical round-off tolerance after an initial search value if located.As each vertex defining a face is read, it is added to an AVL search tree if an equivalent vertex does not already exist in the tree. If the vertex is added, the corresponding vertex reference for the face being added is set to this vertex element in the AVL tree. If an equivalent face exists, the corresponding vertex reference for the face being added is set to the equivalent vertex already in the AVL tree and the vertex data read is discarded. The AVL search tree data structure works well to provide an efficient equivalence detection mechanism for vertices being added.

The AVL search tree data structure works well to provide an efficient equivalence detection mechanism for vertices being added. After vertex merging is completed, face adjacency relationships must be determined. 5.2 Searching for Face AdjacenciesFace comparisons are performed in an attempt to locate face adjacencies. Recall that face adjacencies are determined for each face in the model by sequentially stepping through a list of model faces, searching for adjacent faces, and cross-connecting references when adjacencies are located. A linear list of faces provides the ability to sequentially process faces, but it is clearly not the solution for an efficient searching structure.256The structure imposed to effect efficient adjacency searching is simply a list of face pointers attached to each vertex in the model.

As new vertices are created or existing vertices are referenced by a given face, the reference value of this face is added to a list of face references for each vertex which defines the face. Consequently, each vertex knows of every face which references it after all model faces have been read from the data file. Searching for adjacencies is greatly simplified using this structure. Instead of searching every face in the model for a face which shares two vertices with a test face, now only the faces contained in the face lists corresponding to the three vertices defining a test face need to be searched.for each edge of a test face, all other model faces in a list structure were searched to determine adjacency relationships, this would be of order O(3n2) where n denotes the number of facets in the model.

However, by maintaining for each vertex a list of faces which share it, searching complexity is reduced to an order O(3nk2) problem where n is again the number of facets in the model and k is the average number of faces sharing each model vertex. Figure 5.1 depicts the data available for a given face and is useful in understanding how adjacencies are determined. V 1.",."'".",..",..",."",.",..",."",) List of Face ReferencesFb FnFaF13 FaFeFe Fn.......-..Figure 5.1 - Face Data Used for Adjacency SearchingConsider the searches necessary to locate the facet adjacent to face Fn which shares edge £13. Instead of searching all facets in the model, the search is confined to the lists of face pointers maintained by vertices V1 and V3. The reference to face Fn should appear in the three vertex lists; however, there should be one and only one other identical face reference contained in both the face lists for V1 and V3.

The reference to face Fn should appear in the three vertex lists; however, there should be one and only one other identical face reference contained in both the face lists for V1 and V3. This should reference the face which shares edge £13 with the facet Fn under test and is labeled F13 in the figure. This provides a useful way to bound the search space which must be traversed each time an adjacency must be sought. Since in most cases, the number of facets in a model far exceeds the number of facets sharing a particular model vertex, this provides a significant reduction in the number of elements which must be searched throughout the adjacency determination process. It is also important to note that only face references, not actual vertex coordinate triples, must be tested during searching.

It is also important to note that only face references, not actual vertex coordinate triples, must be tested during searching. This too contributes to the efficiency of this algorithm.6 Results & ConclusionsThere is a definite up-front cost associated with generating topological information from a "bucket of facets" model representation such as the STL de facto standard.257pre-processing adds information which can be used to model processing operations.only a "bucket of facets", can be computationally ........"',..."","'.....", is taken. Appropriate data structures and algorithms gexlen:ltlcm an profitable exercise. An important lesson ... increase in storage requirements can effect a......._ .... 1"'>......... -reference lists for each vertex increases memory "" ...r"T~,r"IO..", a very necessary performance increase.

increase in storage requirements can effect a......._ .... 1"'>......... -reference lists for each vertex increases memory "" ...r"T~,r"IO..", a very necessary performance increase. By using an for model topology a model with approximately 82,500 facets, 8.6 to generate its topology when the brute However, the 0(3nk2) algorithm performed the same machine. It is important to note that these file. Preliminary evaluation of an binary format STL files suggests that the contrast tree approaches is even larger because parse timean 0(3n2) algorithm, processing:r..... "..,,_lJ ..."" ....."-'Jl.Jl""" wereASCII""""'0""""'''CPU minute figure cited above[12].important for successful model processing, and endeavor. Unfortunately, the SFF community's STL defacto models does not support the definition of model topology . ...'Ol",r\1I"II,C>'h........ t-l:)rl from an unordered "bucket of facets" provided by the STL significant searching through model facet data which is part models.

...'Ol",r\1I"II,C>'h........ t-l:)rl from an unordered "bucket of facets" provided by the STL significant searching through model facet data which is part models. The approach presented in this paper magnitude performance improvement when contrast to aisgeIlenlUCln approach.information should be generated during model tessellation , ..... by a CAD system. This would not only prevent the need to it would also make data transfer more robust and concise. The RPI , a significant redundancy reduction while increasing the information ~" by utilizing topological information [9]. While facet models may fade as content the primary representation for future SFF data exchange, topological information and some of the lessons learned from this work will remain important for higher-order geometrical..'VI""'V ...,Jj>.,''''''1-'''''''''...

'-"""Jl.Jl ..VJlJllJaaspatch models.Acknowledgmentsresc~an~n was supported by NSF Grant DDM-8914212 as a subcontract Texas Solid Freeform Fabrication program, the New York State the Office of Naval Research, and other grants of the (RDRC) Industrial Associates Program. Any ""....,Jl.. ""........ lJJl. ...., ...JllJ, or recommendations expressed in this publication are not necessarily reflect the views of the National Science Center for Advanced Technology, the Office of Naval....U ........ lJlWl .......... sponsors. to thank Dick Aubin, Pratt & Whitney (a division of United a number of industrial STL models. A special thanks to u ......·UUJl"" comments and ideas on early drafts of this paper, and tohelpful comments.258References1. "Stereolithography Interface Specification," 3D Systems, Inc., June 1988. 2. Michael E. Mortenson, Geometric Modeling, John Wiley & Sons, Inc., 1985. 3. Stephen J. Rock, "Solid Freeform Fabrication and CAD System Interfacing," M.S.

2. Michael E. Mortenson, Geometric Modeling, John Wiley & Sons, Inc., 1985. 3. Stephen J. Rock, "Solid Freeform Fabrication and CAD System Interfacing," M.S. Thesis, Rensselaer Polytechnic Institute, Troy, NY, Dec. 1991. James V. Miller, "On GDM's: Geometrically Deformed Models for the Extraction of Closed Shapes from Volume Data," M.S. Thesis, Rensselaer Polytechnic Institute, Dec. 1990.4.5. Donald Hearn and M. Pauline Baker, Computer Graphics, Prentice-Hall, Inc., 1986. 6. Kevin Weiler, "Edge-Based Data Structures for Solid Modeling in Curved-Surface Environments," IEEE Computer Graphics and Applications, Vol. 5, Num. 1, pp. 21 40, Jan. 1985.7. Kevin Weiler, "The Radial Edge Structure: A Topological representation for Non Manifold Geometric Boundary Modeling," in: Geometric Modeling for CAD Applications, M. J. Wozny, H. W. McLaughlin, J. L. Encarnacao (eds.), North Holland, pp. 3-36, 1988. Stephen J. Rock and Michael J.

In metal Additive Manufacturing (AM), single tracks are widely used to verify and validate simulation and process models. This work describes the creation of a holistic L-PBF single track data set incorporating a large domain of process conditions for AlSi10Mg. The 352 single tracks are characterized through a series of high throughput methods including white light interferometry, automated microstructure analysis, and in-situ high-frequency (up to 200 kHz) coaxial melt pool monitoring. This data constitutes a large database of process parameters, high- resolution measurements, and geometry information for data driven analysis, including machine learning. In one approach this data is used to correlate track shape and melt pool characteristics to in-situ measurements.IntroductionThe success of the L-PBF process can be distilled to the stability and reliability of themelt pool process.

There is significant research in micro- and mesoscale computational modeling of single and multi track geometry [1, 2]. Coupled with numerical modeling there is also a significant amount of experimental research in high-speed x-ray imaging of the L-PBF process. Existing research has mapped process conditions to single or multi track specimen geometry [3, 4, 5, 6]. This mapping has been done for a limited number of alloys and has hardly, if at all, been conducted with a high level of monitoring data including extensive parameter sweeps. Studying single or multi track level geometry can be used to validate simulation and computational results as well as to evaluate process parameters. Previous studies [7, 8] have included various characterization methods including White Light Interferometry (WLI) and microstructural analysis. The outcome of this type of research is generally a 2D plot mapping process conditions to phenomena such as balling.

The outcome of this type of research is generally a 2D plot mapping process conditions to phenomena such as balling. The dataset presented in this work, however, is significantly richer in completeness and data. This richness will enable research into control schemes that can dynamically improve the stability and reliability of the melt pool process and arguably expand the process window. This data set includes the process parameters, on-axis melt pool monitoring signals, WLI topology scans, and cross sectional images for 352 lines. Most importantly, all these data can be spatially aligned for use in a variety of different research thrusts.1542MethodologyA total of 352 AlSi10Mg tracks were produced for this data set. The tracks were meltedonto 16 additively manufactured sample blocks. The blocks were produced with 30 µm layers and manufacturer-recommended hatch scan speed of 1550 mm/s, hatch laser power of 360 W, and hatch spacing of 150 µm.

The blocks were produced with 30 µm layers and manufacturer-recommended hatch scan speed of 1550 mm/s, hatch laser power of 360 W, and hatch spacing of 150 µm. The tracks themselves were produced from a 30 µm layer, with a range of process parameters across the domain of laser power of 50-400 W and scan speed of 270-2500 mm/s. Gas flow was also varied for some tracks from 300 to 600 L/min. The specific parameters selected are shown in Fig. 1 and are separated into four groups, which are summarized in Table 1. A Latin hypercube (LHC) sampling method was used to select 28 parameter sets spanning the parameter domain for the gas flow LHC group. These tracks were built on one block each at 300, 400, 500, and 600 L/min gas flow rate, for a total of four blocks. The build parameters group consists of 16 parameter sets selected from the hatch and contour parameters used for larger parts. These 16 parameter sets were also used to produce 4-track pads.

The build parameters group consists of 16 parameter sets selected from the hatch and contour parameters used for larger parts. These 16 parameter sets were also used to produce 4-track pads. The third group of parameters consists of 13 parameter sets that were selected as ‘analogous parameters’ where the same linear energy density (0.08, 0.16 or 0.24 J/mm) was achieved through three to five combinations of laser power and scan speed. In Fig. 1, these parameters are connected by lines representing their common energy densities. A separate Latin hypercube sampling generated 35 parameter sets to span the full range of laser parameters for the LHC group. The 64 parameter sets from the build parameters, analogous parameters, and LHC groups were built with gas flow of 600 L/min and spread across a total of 12 sample blocks. Each sample block contained 16 single tracks and four 4-track pads. Two copies of each block were built in the middle region of the build plate.

Each sample block contained 16 single tracks and four 4-track pads. Two copies of each block were built in the middle region of the build plate. One block, which consists mostly of the analogous parameters, was additionally copied to each of the four corners of the build plate to capture potential spatial variations.Figure 1. Power-velocity process map showing the 92 different parameter sets used for the tracks, identified by the group to which each set belongs. The analogous parameters are connected by lines representing their common energy densities. The LHC groups were selected by Latin hypercube sampling.1543Table 1. Summary of parameter groups.

The system monitors the spectral emission of the melt pool at a sampling rate of up to 200 kHz. It collects data from three photodiodes, representing a high-wavelength signal, a low-wavelength signal, and a wide spectrum signal. This data is processed to produce Thermal Energy Density (TEDTM, wide spectrum intensity) and Thermal Emission Planck (TEPTM, ratio of high- and low- wavelength intensities) values. TED and TEP are representative of the total emission from the melt pool and a two color pyrometry measurement calibrated to temperature, respectively [9]. The MPM data for each track was extracted from the whole-build data set.The surface topology of each track is measured on an Alicona InfiniteFocus system (IFMG4f measurement device) at 10x magnification. Each sample block was imaged individually or in a small group, then sections of the full scan were selected for each track.

Each sample block was imaged individually or in a small group, then sections of the full scan were selected for each track. The resulting images had horizontal resolution of approximately 3.5 µm and vertical resolution of approximately 177 nm. An example of this data for a whole block is shown in Fig. 2a, and for a single track from that block in Fig. 2b.The cross section images of each track were acquired by cutting and polishing the sampleblocks after they were removed from the build plate. Each block was sectioned once, in an arbitrary location roughly in the center of the tracks. The blocks are sectioned on a Stuers Accutom-5 sectioning saw and hot mounted in Polyfast. The samples were polished with a Buehler Automet 250 autopolisher and polished to 0.06 microns with a colloidal silica suspension. The sample thicknesses were measured at each step to enable estimation of the polished plane relative to the topological and MPM data.

The sample thicknesses were measured at each step to enable estimation of the polished plane relative to the topological and MPM data. A Nikon MA200 microscope was used at a magnification of 10x. Images were again collected for an entire block of 20-28 tracks and trimmed to a single image for each track. An example of this cross section is shown in Fig. 2c.1544Figure 2. Examples of topology data for (a) a whole block and (b) a single track with the scan direction from the bottom up, with track length of approximately 6 mm, and (c) optical cross section image.AnalysisThe methodologies above created three sets of data for each single track: MPM data,representing both the calibrated temperature measurement and the total emission measurement along with laser location information; the surface topology of the track; and cross section image of the track. The further analysis of these data types was partially automated to expedite analysis.

The further analysis of these data types was partially automated to expedite analysis. Full automation was not feasible because of the noisy nature of the topology and cross section imaging of single tracks on additively manufactured surfaces, and the wide range of parameters which lead to a variety of track shapes.Initial processing on the topology data was conducted by manually labeling tracks as either ‘complete’, ‘partial’ or ‘missing’, representing the cases of a coherent track, a broken up or segmented track, or no track, respectively. These cases are demonstrated in Fig. 3. The topology data for each track was independently labeled by four people for a total of five times. Final labels were determined by a majority of the five labels. The start and end positions of each track were manually identified. The cross section images were manually annotated in ImageJ, with the region of the image identified as the track’s cross section marked as white.

The cross section images were manually annotated in ImageJ, with the region of the image identified as the track’s cross section marked as white. The cross sections were not etched, therefore the track region was not always clear, depending on polishing and microscope settings. Because of this, not all images could be annotated. Further analysis shown here was limited to single tracks labeled as ‘complete’ and with successful cross section annotations, though valuable information could also be gleaned from the incomplete tracks. A1545few additional tracks were removed for various processing issues. This resulted in a total of 156 single tracks down selected for further analysis.Figure 3. Example topolgy data for (a) complete track, (b) partial track, and (c) missing track. These tracks are approximately 8 mm long.For the down-selected 156 tracks, the melt pool cross section was analyzed by calculatingthe dilution and aspect ratios from the annotations.

These tracks are approximately 8 mm long.For the down-selected 156 tracks, the melt pool cross section was analyzed by calculatingthe dilution and aspect ratios from the annotations. Here, dilution ratio is defined as the ratio between the depth of the track below the level of the substrate and the height of the track above the substrate. Aspect ratio is defined as the ratio between the height of the track above the substrate and the width of the track.Figure 4. An example of the topology analysis, taking a slice of the topology data perpendicular to the track and extracting the peak, left, and right edge for the slice.1546The height and width for the full length of each track was calculated from the topologydata. The topology data was sliced perpendicular to the track, and the height, left edge and right edge were extracted for each slice based on the track profile cross section at that point. These points are shown in an example in Fig. 4.

These points are shown in an example in Fig. 4. The width was then calculated from the difference between the left and right edges. The MPM data, height, and width data were then processed so that each data set had the same spatial resolution along the track. This registration was conducted by generating 1000 y-positions evenly spaced along the track, then selecting the height, width, and MPM data values with y-positions closest to each of the 1000 generated points.ResultsThe track cohesion classification of the 352 tracks shows a general trend of both lack offusion (LOF) and balling defects. In Fig. 5, as the energy density of the tracks decreases, the cohesion of the tracks decreases, shifting from complete track, to incomplete, to absent. Additionally, as the scan speed increases, even with a corresponding increase in laser power, the tracks become less cohesive, demonstrating a balling up effect.

Additionally, as the scan speed increases, even with a corresponding increase in laser power, the tracks become less cohesive, demonstrating a balling up effect. Tracks in the balling and LOF regions were excluded from further analysis here because the inconsistent topology of the tracks requires a different processing technique to generate useful information.Figure 5. Power-velocity process map showing the cohesion of 352 tracks across 92 parameters. Due to the repetition of parameters across duplicate sample blocks, some parameter sets have multiple labels.Examination of the topology data with the MPM data along the length of the track showssome correlation. Comparison of the average MPM data values for each track with the cross section data and line cohesion classifications did not demonstrate a clear relationship. The reason for this contrast in results is likely the loss of data resolution through the averaging.

The reason for this contrast in results is likely the loss of data resolution through the averaging. Since this averaging is not present with the topology/MPM data comparison, a clearer pattern is able to emerge. Fig. 6(a) shows this pattern for the height of the track and the TED signal. Fig. 6(b)1547shows the relationship between the width of the track and the TED signal. For both of these plots, the TED signal was smoothed with a moving average of 10 samples. These plots demonstrate that for some regions of the track, the TED signal and track dimensions appear to be moving in similar patterns. However, when applied to the aggregate data of the 155 single tracks selected for this analysis, there does not appear to be a prevailing statistical correlation between MPM data values and track dimensions. This suggests that the relationship between the MPM data and the track dimensions is more complex.

This suggests that the relationship between the MPM data and the track dimensions is more complex. The more complex relationship is consistent with the nature of the differences in the data sets. The track dimensions are related to the thermal gradient and morphology of the melt pool at multiple points in time, while the MPM data is only able to capture some of these characteristics at a single point in time. It is also possible that the MPM data shown here captures only one aspect of the changes in melt pool characteristics that leads to the variations in track dimensions and doesn’t capture variation due to, for example, variation in the substrate height.Figure 6. The TED MPM signal and the (a) height and (b) width of the resultant trackseem to move in similar patterns for certain sections of the track.1548Conclusions and Future WorkSingle tracks are the most fundamental component of the L-PBF process.

Since L-PBF isan extremely localized manufacturing process, understanding the behavior of single tracks with relation to melt pool monitoring data is crucial to improved process understanding and control. The work presented here demonstrates a framework to generate and analyze large sets of single tracks. The generated data of over 20 GB shows track morphology consistent with expected lack of fusion and balling behavior and illustrates a potential connection between MPM data and resultant track shape. These demonstrations show that the generated data set has the potential to provide useful insights to connect melt pool monitoring to observed track shapes. Further analysis is needed to fully characterize these relationships, including a more robust model to connect complex MPM data behavior to track morphology. Additional analysis of the tracks themselves also includes sample etching, porosity calculations, and analysis of incomplete tracks.

Additional analysis of the tracks themselves also includes sample etching, porosity calculations, and analysis of incomplete tracks. The data set generated here is available at: https://github.com/ConorPorter3/Single- Tracks.AcknowledgementsResearch was sponsored by the Army Research Laboratory and was accomplished under Cooperative Agreement Number W911NF-20-2-0292 and W911NF-21-2-02199. The views and conclusions contained in this document are those of the authors and should not be interpreted as representing the official policies, either expressed or implied, of the Army Research Office or the U.S. Government. The U.S.

Many researchers employ FEM to analyze the effects of different filling strategies. However, they mainly majored in regular simple path (e.g. line and circle).This paper presents the method to generate meshes in deposition region of complicated path which can be used in FEM. First, the contour offset path and skeleton path is introduced. Then, the deposition region of each path is created by offsetting the paths inward and outward. Afterwards, quadrilateral meshes are constructed within each region. Deposition region is approximated by meshes.Keyword: deposition path, deposition region, quadrilateral meshesIntroductionAdditive Manufacturing (AM), or Rapid Prototyping (RP), has rapidly been developed in recent decades. Different from the traditional material subtract manufacturing method, it fabricates part layer by layer. In AM field, Wire and arc based Additive manufacturing has attracted more and more attention of researchers and institutions.

In AM field, Wire and arc based Additive manufacturing has attracted more and more attention of researchers and institutions. Compared to AM with powder as feed stock material [Sing[1], Ding[2], Ding[3]] or laser as energy source [Ding[4]], the wire and arc AM has a lot of advantages, such as high material utilization, simple and inexpensive equipment and high production efficiency[Gu [5], Wang[6]].The filling path plays an important role in AM process. It directly influences the geometric accuracy and mechanical property of fabricated part. There are two major basic filling patterns: raster scanning path and equal distance contour offset path [Ding D[7]]. Raster scanning path generation algorithm is very simple and efficient. But it scans in the same direction that would cause stress concentration leading to warping. Furthermore, the tool needs to cross the cavity frequently for part with holes. And the contour has a low accuracy because the raster path starts and ends in contour.

Furthermore, the tool needs to cross the cavity frequently for part with holes. And the contour has a low accuracy because the raster path starts and ends in contour. Equal distance offset path generation algorithm iteratively offsets the outer contour of layer polygon by a fixed distance until filling the whole interior domain. This method change filling direction gradually (but still same in segment part) in each path which972can decrease stress concentration. Contrast to raster path, the offset path is continuous in contour that it can improve the surface accuracy. However, the contour offset algorithm is very complex and time consuming. G.Q. Jin [8] combines the two type of path. In the contour, the contour offset path is adopted to ensure the accuracy, and the zigzag raster path with the best slope degree is employed in the interior to achieve minimum manufacture time.

In the contour, the contour offset path is adopted to ensure the accuracy, and the zigzag raster path with the best slope degree is employed in the interior to achieve minimum manufacture time. Y Yang [9] gets the contour offset path through the interior recognition and complex self-intersection and sharp corner treatment. DS Kim [10] generates contour offset path based on Voronoi Diagram which do not needs self-intersection treatment. Y.Ding [Ding [11]] mapped the revolved overhanging structures to a plane base and filled the contours by offsetting the medial axis until the whole contour is covered. This method initially provided a solution for AM building overhanging structures in 8-axis system.Apart from the pattern of filling path, the deposition sequence determined by the paths also affects the part property in fabrication process. Sattari-Far I [12] studied the welding distortions in pipe-pipe joints of AISI stainless-steel type caused by different deposition sequence. A.H.

Sattari-Far I [12] studied the welding distortions in pipe-pipe joints of AISI stainless-steel type caused by different deposition sequence. A.H. Nickel [13] investigated two different deposition direction of raster pattern and two different deposition sequence of offset pattern. The FE analysis of four deposition sequence of multi-pass single-layer was performed by M P Mughal [14]. In the thermal analysis of welding process using FEM analysis software (e.g. ANSYS), the “birth and death” method is a key point because the activate order of elements represents the real deposition sequence. In many researches, only a few simple paths, such as segment, rectangle and circle, are considered and the activate order is decided manually. However, it is hard to do so for complex paths and elements.Generation of deposition regionIn production of complex thin wall parts, the wire arc welding additive manufacturing (WAAM) improve manufacturing technology can substantially efficiency.

Contour offset paths, a common type of deposition path used in WAAM, are simple polygons derived from slice contours by inward offsetting. The method proposed in this study uses Voronoi diagram to create a contour offset path. For a given set of points and edges in a plane, a Voronoi diagram partitions the plane into regions, with each region including one point or one edge in its interior [Aurenhammer [15]]. These points and edges are called point sites or edge sites, and the regions are called Voronoi cells. For any point in a cell, its distance to the site associated with it is shorter than its distance to any other site. The boundaries of the Voronoi cells, i.e. the lines separating the regions, are referred to as Voronoi edges.

The boundaries of the Voronoi cells, i.e. the lines separating the regions, are referred to as Voronoi edges. For planar slice contours, sites are vertices and edges of polygons, and cell boundaries are made up of vertex angle bisectors, perpendiculars to edges at their endpoints, lines lying equidistant between parallel edges, and parabolas between points and segments. All line segments constituting the inner and outer contours offset within their own Voronoi cells and are joined to each other at their endpoints. This avoids dealing with973contour intersections and self-intersections and thereby allows for efficient generation of a contour offset path based on Voronoi diagram. Figure 1 shows arbitrary slice contours. Figures 2 and 3 show the corresponding VD and contour offset path.Fig.1. two slice contoursFig.2. the Voronoi diagram of contours974(a) all deposition paths(b) one of the deposition paths(c) all deposition paths(d) one deposition pathFig.3.

two slice contoursFig.2. the Voronoi diagram of contours974(a) all deposition paths(b) one of the deposition paths(c) all deposition paths(d) one deposition pathFig.3. Deposition paths based on Voronoi diagram (purple outer polygons areslice contours, the red&yellow inner polygons are deposition paths)In addition to contour offset path, skeleton is often extracted from VD as filling paths, especially for a part whose wall thickness is smaller than the width of a single weld bead. The skeleton is a part of the Voronoi diagram. The algorithm proposed in this study provide two types of skeletons: skeleton derived from a single contour (bisectors of the edges in one polygon) and the skeleton derived from different contours (bisectors of edges in two polygons), as shown in Figure 4.975Fig.4.

skeleton (medium purple) and contours (black)In WAAM, filling paths determine the position and sequence of energy input and thus have a great influence on the deformation and stress state of fabricated part. The finite element method (FEM) is usually employed to analyze the distortion, temperature, and stress experienced a part during deposition. This method requires meshing of the real deposition region of the path. Deposition region refers to the ring-shaped region occupied by the material deposited from the melted wire. In this paper, a given deposition path is used as the input data, and the Voronoi algorithm is applied again to offset the path inward and outward separately, in the same manner as in the path generation process. Then two simple polygons are created as the outer and inner contours. The offset distance is half the effective width of a single deposition path, which depends on actual process parameters.

Then two simple polygons are created as the outer and inner contours. The offset distance is half the effective width of a single deposition path, which depends on actual process parameters. Later, a bounding box is used to determine the inclusion relation between the outer and inner contours. The region enclosed by the two contours represents the effective deposition region. The vertices of the outer contour are ordered anti-clockwise, while those of the inner contour are ordered clockwise. Therefore, as the wire moves along the outer contour, the interior domain is on the left-hand side, while as the wire moves along the inner contour, the interior domain is on the right-hand side.976Fig.5. Path (red polygon) and its deposition region contours (purple polygons)As described above, the deposition region of a single path is defined by two polygons whose shapes are similar to the path’s.

Path (red polygon) and its deposition region contours (purple polygons)As described above, the deposition region of a single path is defined by two polygons whose shapes are similar to the path’s. The size of cells within the deposition region can be modified by adjusting the lengths of line segments during discretization of the region. In the method proposed in this study, offset polygons (including paths and the contours defining the deposition region) are constructed from line segments and arcs based on VD. The size of cells can be adjusted by discretization of line segments and arcs and specific parameters can be input by users.The proposed method does not directly discretize the deposition region. Instead, it discretizes the original path before offsetting so that discrete deposition contours can be naturally obtained after offsetting the path inward and outward. In addition, the arcs of the path need to be discretized into line segments.

In addition, the arcs of the path need to be discretized into line segments. The purpose is to convert the path into a simpler polygon and thereby facilitate VD construction and path offsetting. Line segments can be discretized into several short segments with a given length or into a preset number of segments. Arcs can be discretized into segments with a given arc length or central angle. The vertices in the final discrete path are classified into three types: original vertices (0), points on line segments between their endpoints (1), and points on arcs between their endpoints (2). These vertices are denoted by V0, V1, or V2, with the integer identifiers (0, 1, and 2) indicating their attributes, i.e. types, as shown in the figure below.011102220Fig.6.

These vertices are denoted by V0, V1, or V2, with the integer identifiers (0, 1, and 2) indicating their attributes, i.e. types, as shown in the figure below.011102220Fig.6. Types of points (0 - end points of line segment or arc, 1 - points in linesegment, 2 - points in arc)977After discretization of line segments and arcs, the discrete points on line segments between their endpoints automatically offset, generating discrete points on the line segments constituting the deposition contours. Most of the discrete points on the inner contour are paired with those on the outer contour. Then any two pairs of points on the contours can form a quadrangle. The resulting discrete points on the offset contours have the same attributes as their source points on the original path. These points are denoted by V0 or V1.In the VD offset method, an arc will be generated when a convex vertex of a polygon offsets outward or a concave vertex offsets inward.

These points are denoted by V0 or V1.In the VD offset method, an arc will be generated when a convex vertex of a polygon offsets outward or a concave vertex offsets inward. In this study, an offset arc can be generated by offsetting a vertex V0 or V2. An offset arc generated from a vertex V0 needs to be discretized; the length of the resulting segments depends on the discrete length or central angle of the arc. The two endpoints of the offset arc are designated V0 while other discrete points on it are designated V2. If an offset arc is derived from a vertex V2, it is just approximated by a line segment connecting its two endpoints, which are denoted by V2. As mentioned above, vertices V2 on the path are continuous and form arcs. Offsetting these arcs also produces arcs approximated by discrete segments. However, it is necessary to further discretize these arcs in order to obtain cells with desired sizes.

Offsetting these arcs also produces arcs approximated by discrete segments. However, it is necessary to further discretize these arcs in order to obtain cells with desired sizes. Since these offset arcs are concentric with their source arcs on the path, the radius of each offset arc is its source arc’s radius plus the corresponding offset distance. The further discretization is also based on a preset arc length or central angleThe discretization of original path enables automatic generation of discrete points on the contours during offsetting, thereby reducing the amount of calculation. Most quadrilateral cells in the mesh generated are obtained by combining triangles without complex calculation.Fig.7. Discrete contours978Quadrilateral mesh in deposition regionThis section describes how to generate a finite element mesh in the effective deposition region. A mesh normally comprises quadrangles, which are derived from triangles, and a small number of triangles.

A mesh normally comprises quadrangles, which are derived from triangles, and a small number of triangles. Based on the discrete polygons (offset contours), a Delaunay triangulation is constructed to mesh the polygons. The triangles in the triangulation share vertices with the polygons, as shown in the figure below.Fig.8.

The triangles in the triangulation share vertices with the polygons, as shown in the figure below.Fig.8. Triangle meshesBased on attributes of vertices, the triangles in a triangle mesh can be divided intoseveral types as follows:(1) For a triangle having vertices V1, V1, and V1, it is classified as T0 if the three vertices lie on two contours; if they are distributed on the same contour, it is classified as T4;(2) For a triangle having vertices V2, V2, and V2, it is classified as T1 if itsvertices lie on two contours; if they are on the same contour, the triangle type is T5;(3) For a triangle having vertices V0, V2, and V2, its type is T2 if its vertices aredistributed on two contours; otherwise, its type is T6;(4) Other triangles have their vertices coinciding with the vertices of the contours and link two line segments, or a line segment and an arc.

These triangles are called link triangles and denoted by T3Triangles T0, T1, T2, and T3 all have only two adjacent angles.97903230Fig.9. triangle typesNext, triangles can be combined into quadrangles using the method below: (1) Randomly select a triangle with types T0, T1, and T2 as the start triangle.Then set the start triangle as _current and set quadrilateral index, _qidx, as 0;(2) Traverse the adjacent triangles of _current, and store triangles T0, T1, T2, andT3 that have not been visited in the candidate triangle container, _candidates;(3) If _candidates is empty, go to (8); (4) If the attribute of _current is T2, set _current as degenerate quadrangle and store it in the mesh container. Then assign the quadrangle’s index to corresponding triangles.Set the current triangle as the combine triangle and go to (7);(5) Each candidate triangle can combine with _current into a quadrangle.

Then assign the quadrangle’s index to corresponding triangles.Set the current triangle as the combine triangle and go to (7);(5) Each candidate triangle can combine with _current into a quadrangle. Find the quadrangle whose interior angles have the minimum standard error and then set the corresponding candidate triangle as _combine;(6) If the attribute of _combine is different from that of _current, go to (4);otherwise, store this quadrangle in the mesh container;(7) Traverse the adjacent triangles of _ combine and set a triangle T0, T1, T2, orT3 that has not been visited as the new _current. After that, go to (2); and(8) Traverse unhandled triangles and find its adjacent quadrilateral. If the quadrilateral is a triangle (degenerate), directly merge it with this degenerated quadrilateral.

If it is real quadrilateral, split the quadrilateral into a quadrangle and a triangle and then combine this two triangle (unhandled one and split one).980STARTIf type of current triangle is 2classified trianglesYESNOstore current in meshesfor each triangle in candidatesfor each triangle in trianglesNOset combine triangle to be currentselect a suitable triangle as a combine triangleIf the type of unvisited triangle if 0,1,or 2YEScurrent = triangleIf type of combine triangle is equal to type of current triangleYESfor each unvisited adjacent triangle of currentmerge combine and current to a quadrilateral, and store in the meshesIf the type of triangle is 0,1,2 or 3for each unvisited adjacent triangle of combine triangleNOstore triangle in candidatesIf the type of triangle is 0,1,2 or 3NOYESIf candidates is emptyYESset the triangle as new current trianglefor each unvisited trianglesMerge the triangle into its adjacent quadrilateralENDFig.10.

Flowchart of generating meshesFig.11. Quadrilateral meshes (blue) and pathAt last, the mesh cells are ordered based on sequence of points in the path, in981order to ensure that the sequence of the cells is consistent with the deposition sequence. Therefore, in the FEM analysis, the activation sequence of elements can be determined easily without any calculation.ConclusionIn this study, the contour offset path and skeleton path are created for the fabrication of thin wall part using WAAM technology. On the purpose of learning the deformation and stress distribution of part caused by the deposition path, the method of meshing in the real deposition region is introduced. The meshes in the deposition region mainly consist of quadrilateral meshes and also there exists little triangles sometimes.In the future work, the optimization of deposition path based on the FEM analysisis the primary aim.References[1] Sing, Swee Leong, et al.

"Laser and electron‐beam powder‐bed additive manufacturing of metallic implants: A review on processes, materials and designs." Journal of Orthopaedic Research 34.3 (2016): 369-385.[2] Ding, Yaoyu, James Warton, and Radovan Kovacevic. "Development of sensing and control system for robotized laser-based direct metal addition system." Additive Manufacturing 10 (2016): 24-35.[3] Ding, Yaoyu, and Radovan Kovacevic. "Feasibility study on 3-d printing of laser-based metal additivemetallic structural materials with manufacturing." JOM 68.7 (2016): 1774-1779.robotized[4] Akbari, M., Ding, Y., & Kovacevic, R. (2017, June). Process Development for a Robotized Laser Wire Additive Manufacturing. In ASME 2017 12th International Manufacturing Science and Engineering Conference collocated with the JSME/ASME 2017 6th International Conference on Materials and Processing (pp. V002T01A015-V002T01A015). American Society of Mechanical Engineers.

NERATION OF MICRO MECHANICAL DEVICES USING STEREO LITHOGRAPHY A. Neumeister*, R. Himmelhuber+, T. Temme* and U. Stute* * Laser Zentrum Hannover e.V., Hollerithallee 8, 30419 Hannover + micro resist technology GmbH, Koepenicker Str. 325, 12555 Berlin Reviewed, accepted September 14, 2006 Abstract A high resolution machining setup for creating three-dimensional precision components from a UV-curable photo-resin has been developed. By using frequency-converted diode-pumped solid state lasers, functional micro-mechanical devices are directly fabricated in a successive layer-by- layer fashion. Within the direct generation of micro assemblies having moving components without further assembly of parts will be presented. The micro system design is based on user-defined 3D-CAD data and will completively be built up within the fabrication cycle.

The micro system design is based on user-defined 3D-CAD data and will completively be built up within the fabrication cycle. By using specially developed μSL materials with suitable properties for micro- mechanical parts, the development from Rapid Prototyping towards Rapid Production of small series is intended.this paper,Introduction Nowadays, micro system technology is a rapidly expanding market with applications in both research and industrial sector in the field of automotive, medicine or biotechnology [1]. Surfaces of artificial lenses for the human eye made of a hydrophilic acrylate can be mechanically machined with a roughness in the micron range [2]. The trend of miniaturisation and integration of functional micro systems is becoming more important and requires the further development of new and existing production technologies. As consequence of scaling down the part size, the integration of micro components becomes more demanding.

As consequence of scaling down the part size, the integration of micro components becomes more demanding. For the extensive handling of micron sized parts, a high degree of automisation is intended to meet the distinctive quality requirements in micro technology [3]. In order to increase the level of product innovation, the use of materials with specific properties in the optical or mechanical range is essential. The combination of both advanced material properties and innovative process technology can further lead to products with high complexity, leading to new functionality at lower costs. Within this paper, the authors present a free form micro fabrication technology for the production of three dimensional complex objects based on the stereo lithography (SL) principle.

Within this paper, the authors present a free form micro fabrication technology for the production of three dimensional complex objects based on the stereo lithography (SL) principle. Basically, SL was used to generate prototypes to provide important visual and haptic data for parts that will be subsequently fabricated in mass production processes like injection molding or die casting (Rapid Prototyping). Recently, layer-based fabrication techniques are applied to directly build up functional parts with properties close to materials that are used in batch production (Rapid Manufacturing). The advantages of such fabrication processes are low production costs for small batch sizes, high flexibility in the product design and its change and the fast delivery of functional parts. The usability of this production technology for special applications in the micro technology will be demonstrated during this presented work.

The usability of this production technology for special applications in the micro technology will be demonstrated during this presented work. Main innovation of the developed technique is the direct production of mechanical assemblies with movable mechanical components without the necessity of conjoining individual parts. Since no additional assembly steps are required during the production cycle, the developed technology enables the cost-effective fabrication of a small number of high-complex systems. Innovative resin materials with material properties after curing close to some ceramics are used for production. The curing properties can be adjusted towards lateral and vertical dimensions for high resolution processing.12Recently, the suitability of custom made stereo lithography systems for the generation of micro mechanical parts has been shown for the fabrication of individual precision parts [4-6].

By using an Argon-Ion laser operating at 351.1 nm, gearwheels with diameters in the range of 500 – 700 μm have been produced. The geometrical information was delivered by a spatial light modulator as a dynamic lithographic mask. The material which was used was the Ciba-Geigy Cibatool SL 5180 leading to a layer thickness in-process of around 50 μm. The number of layers was up to 105. Micro gears with a diameter of 1500 μm have been produced by applying a UV spot curing system which is equipped with a broadband high pressure mercury lamp (300 - 470 nm). The lateral resolution that has been achieved during the experiments was down to 20 μm. In order to use resin materials with high viscosity, ultra-sonic vibration was applied to reduce the polymer settling time.

In order to use resin materials with high viscosity, ultra-sonic vibration was applied to reduce the polymer settling time. To further demonstrate the capability of micro SL for complex structuring, the fabrication of functional micro assemblies with moving parts will be in the following presented.Experimental set-up The principle set up of the applied micro stereo lithography machine is depicted within figure 1.quasi-cw laser pulses @ 355 nm 20 mW, 10 ps, 100 MHzgalvano scanner + scanning lensresin deposition system with roller coatersubstrate platformcompensation platformresin processing chambersubstrateresin levelFigure 1:Principle setup of the processing chamber of the applied micro stereo lithography setupThe custom micro machining system is set up on an optical table that carries the optical setup (laser source with beam guidance) and a granite based frame with the resin processing chamber, positioning stages and laser scanning system (Figure 2).

The system is equipped with in total four piezo controlled positioning stages offering a maximum encoder based resolution of 300 nm. The substrate displacement inside the resin processing chamber is done in negative z-direction below the resins surface. The substrate is mounted on a custom made stainless steel frame. In order to keep the resin surface at a constant level during the positioning cycle, a frame identical in construction to the substrate frame can be moved in the opposite direction. The coating mechanism consists of a novel roller coater that is in constant contact with the resin. The13coater laminates the substrate with a resin layer of defined thickness. The focal position can be precisely adjusted by moving the scanner system in the vertical direction with a micro meter screw offering a 1 μm resolution at a maximum travel of 50 mm. For illumination purpose, a frequency tripled diode pumped solid state Nd:YAG oscillator (355 nm) is used.

For illumination purpose, a frequency tripled diode pumped solid state Nd:YAG oscillator (355 nm) is used. Pulse trains of 10 pico seconds (ps) at a repetition rate of 100 MHz are provided. The calculated focal radius is around ωf = 7.5 μm. The focused laser beam is scanned on the photo resins surface with two perpendicular oriented galvanometric mirrors of a scanning system with a working field of AW = 50 x 50 mm².Figure 2:Picture of the custom made micro stereo lithography set-up consisting of granite based frame and the optical setup with AOM and laser source (b) and beam guidance (c)Experimental procedure For the fabrication of three dimensional precision parts for micro mechanical applications, a specially designed inorganic-organic UV-curable hybride polymer from the type Ormocer® b59 has been used. In order to apply a suitable material for the micro stereo lithography process, the basic material has been chemically modified.

In order to apply a suitable material for the micro stereo lithography process, the basic material has been chemically modified. Afterwards, a set of material samples has been prepared and characterised in a first step. Main aspect of investigation during the characterisation process has been both the lateral and vertical curing resolution and the viscosity of the resin. In the following, the chemical modification of the basic material with additives for the manipulation of both the curing properties and viscosity will be explained.Material properties, preparation and modification: The hybrid polymer was synthesised in a water free sol gel process where Diphenylsilandiol reacts with Methacryloxipropyltrimethoxysilane (MEMO) (see Fig. 3) in alkaline conditions. During the sol-gel reaction an inorganic Si-O-Si network is formed [7]. Diphenylsilanediol acts as a network modifier, whether MEMO provides the organic cross-linking functionalities [8].

During the sol-gel reaction an inorganic Si-O-Si network is formed [7]. Diphenylsilanediol acts as a network modifier, whether MEMO provides the organic cross-linking functionalities [8]. The size of the formed oligomers with an inorganic core and an organic shell is typically between 2 and 10 nm [9], but the detailed structure has not been shown yet. After the synthesis and removal of solvents, 1% of a radical initiator (Irgacure 369 from Ciba, used as received) was added. When the initiator is illuminated with UV-radiation it decomposites to radicals, which can start a14radicalic crosslinking reaction between the oligomers.

When the initiator is illuminated with UV-radiation it decomposites to radicals, which can start a14radicalic crosslinking reaction between the oligomers. Those will form a three dimensional network with a high mechanical stability.abOHOSi OHOSiOOOFigure 3:Chemical structure of Diphenylsilanediol (a) and Methacryloxipropyltrimethoxysilane (MEMO) (b)In order to reduce the optical penetration depth of the laser radiation, experiments with several absorbing dye materials (ADM) were carried out. The dyes are supposed to absorb at the wavelength of the laser (355 nm). In Figure 4, the molar absorption coefficient of three potential dye materials can be seen over the wavelength.

The dyes are supposed to absorb at the wavelength of the laser (355 nm). In Figure 4, the molar absorption coefficient of three potential dye materials can be seen over the wavelength. ADM 1 shows the highest absorption coefficient at the laser wavelength.4,01 -m c 1 -lo m3,53,0h t g n e e v a wlr e s aADM1 ADM2 ADM3l/t n e c i f f e o c n o i t b r o s b ai2,52,01,51,0r a o ml0,50,0300320340360380400420440wavelength λ / nmFigure 4:Absorption coeffitient of the dyes over the wavelength.However, the total quantity of the selected dyes that can be mixed into the basic material had to be taken in account. With 0.2 % of ADM1, 0.3 % of ADM2 and 0.5 % of ADM3, the maximum amount of the dyes that could be solved, have been introduced into the inorganic-organic hybrid material.

With 0.2 % of ADM1, 0.3 % of ADM2 and 0.5 % of ADM3, the maximum amount of the dyes that could be solved, have been introduced into the inorganic-organic hybrid material. As a result of the variation in solubility of the three absorbing dyes, the effective resulting absorbance differs from the absorbance of individual dyes.

As a result of the variation in solubility of the three absorbing dyes, the effective resulting absorbance differs from the absorbance of individual dyes. Figure 5 shows the resulting absorbance of the dyes in combination with the hybrid material.As ADM3 shows the highest resulting absorbance at the selected laser wavelength, the further investigations were carried out with mixtures of the hybrid material and ADM3.15300250h t g n e e v a wlADM1 ADM2 ADM31 -m cr e s al/200t n e c i f f e o ci150n o i t b r o s b a100500 300320340360380400420440wavelength λ / nmFigure 5:Resulting absorbance of the dyes in the hybrid materialFurther chemical modifications of the hybrid material have been made to reduce the viscosity since it was know from recent investigations that the high viscosity of the hybrid material leads to a long settling periods after the coating procedure.

Several samples with a low viscosity multifunctional acrylate monomer (MAM) were prepared. The influence of the monomer concentration on the viscosity of the resulting material is shown in Fig. 6.3,53,02,5a P M2,0/y t i s o c s vi1,51,00,501020304050weight percent of monomerFigure 6:Influence of the multi functional acrylate monomer (MAM) concentration on the viscosity of the resulting materialThe modified materials have been filtered down to 0.8 μm before the curing investigations.Basic investigations: The curing properties of the developed material samples have been studied in 2D (irradiation of spin coated samples) and 2 ½ D (irradiation of small resin reservoirs) tests.

For the determination of the lateral resolution cw, rectangular glass samples have been spin coated with a constant layer thickness of 10 μm and subsequently irradiated with lattice structures under the influence of varying laser scanning speed vs at a constant laser power of PL = 300 μ W. (Figure 7). The calculated laser fluence was in the range from H0 = 10 mJ/cm² to 1000 mJ/cm²16and the distance between individual lines of the lattice structure has been 250 μm.

(Figure 7). The calculated laser fluence was in the range from H0 = 10 mJ/cm² to 1000 mJ/cm²16and the distance between individual lines of the lattice structure has been 250 μm. The prepared samples have been cleaned after illumination with a solvent of 2-Propanole and 4-Methyle 2- Pentanone and observed by optical, confocal and scanning electron microscopy (SEM).afocused UV lightbPMMA reservoirspin coatingcuring width cwquartz samplelattice structures with hatch distance dhcuring depth cdFigure 7:Experimental technique for the determination of the curing width cw (a) and curing depth cd (b)The curing width cw for Ormocer b59 with varying parts of ADM3 (0%-0.5%) is depicted within figure 8. According to the graph, the lateral curing width cw strongly depends on the part of the absorbing dye that has been introduced in the basic material. By applying a laser fluence of 500 mJ/cm², a curing width of 70 μm is achieved without the use of an absorbing medium.

By applying a laser fluence of 500 mJ/cm², a curing width of 70 μm is achieved without the use of an absorbing medium. With 0.5 % of ADM3, the curing width is reduced by a factor 1.4 towards 50 μm. Generally, a higher process resolution can be achieved by increasing the part of the absorbing dye. However, it has to be taken in account that the usage of a laser fluence H0 < 100 mJ/cm² might result in an unstable process due to the high gradient.10080m μ /wc60htd w g n i r u ci4020Ormocer b59 ADM 3 0% Ormocer b59 ADM 3 0,1% Ormocer b59 ADM 3 0,25% Ormocer b59 ADM 3 0,5%02004006008001000laser fluence H0/ mJ cm-2Figure 8:The curing width cw versus the laser fluence H0 of modified Ormocer b59 with varying parts of ADM317The smallest line structures that adhere at the quartz material after chemical cleaning have been around 7.5 μm at a laser fluence of H0 = 20 mJ/cm².

Nevertheless, structures could be produced with H0 < 20 mJ/cm², but not measured. The curing width cw for Ormocer b59 with a constant content of 0.5 % of ADM3 and varying parts of MAM (0 %-50 %) is depicted within figure 9.80m μ60/wch t d w g n i r u ci4020Ormocer b59 ADM 3 0,5% Ormocer b59 ADM 3 0,5% MAM 10% Ormocer b59 ADM 3 0,5% MAM 50%002004006008001000laser fluence H 0/ mJ cm-2Figure 9:The curing width cw versus the laser fluence H0 of modified Ormocer b59 with 0.5 % of ADM3 and varying parts of MAM.According to the graph, the lateral curing behaviour also depends on the part of the introduced MAM. With increased content of MAM, the curing width increases. The lateral process resolution can be adjusted within the range 7 μm <cw< 80 μm by modifying the basic material with varying contents of ADM3 and MAM. Here, the highest process resolution is achieved with Ormocer ADM3 (0.5 %) and MAM (50 %).

Here, the highest process resolution is achieved with Ormocer ADM3 (0.5 %) and MAM (50 %). Useful processing dimensions for μSL are achieved in the range H0 < 100 mJ/cm². For the determination of the vertical resolution (curing depth cd) a small test cell according to figure 7b has been set up. By using a rectangular solid of PMMA (5x5x1 cm³) with a blind hole, a resin reservoir has been constructed. The hole has been filled with the resin material and covered with a circular quartz substrate that was in direct contact to the resin material. By aligning the laser focus on the polymer quartz contact plane, structures that adhere at the bottom side of the quartz substrate have been fabricated. In contrast to 2D testing, lattice structures with varying lattice constants have been produced for the determination of the vertical process resolution. The lattice constant (= hatch distance dh) has been reduced until it was below the order of the previously determined curing width cw.

The lattice constant (= hatch distance dh) has been reduced until it was below the order of the previously determined curing width cw. The aim of the investigation was to find out, to what extent the curing depth cd changes, in the condition that individually cured lines of the generated lattice structure overlap. Curing depth measurements have been performed from H0 = 10 mJ/cm² to 200 mJ/cm².

Without parts of ADM3, the curing depth cd of Ormocer b59 is the range 230 μm < cw < 375 μm for dh = 0.001 mm and shows no suitability for the μSL process. The introduction of ADM3 (0.5 %) leads to the effective increase of the achieved vertical process resolution down to 18 μm at dh= 0.001 mm, H0 = 20 J/cm² and 8 μm (dh = 0.005 mm, H0 = 40 J/cm²). Suitable fabrication parameters for layer sizes around 10 μm are in the area of 20 J/cm² < H0 < 200 J/cm² at a hatch distance of dh = 0.005 mm. The hatch distance dh is a degree for the theoretical overlap δ of individual fabricated lines of the lattice structure and can directly be calculated from the focal radius ωf and the hatch distance dh according to δ = (2ωf - dh)/ 2ωf. For dh = 0.001 mm, the overlap is calculated to δ = 93 %, for dh = 0.005 mm the overlap is equal to δ = 66 %. It has been found, that a higher overlap factor leads to increased curing depth cd at identical laser fluence.

It has been found, that a higher overlap factor leads to increased curing depth cd at identical laser fluence. At a laser fluence of H0 = 1 kJ/cm², the curing depth cd = 220 μm for a hatch distance of dh = 0,005 mm. In contrast to this result, a curing depth of cd = 300 μm has been achieved at hatch distance of dh = 0,001 mm. A strong correlation between the overlap of the individual scanned lines and the curing depth at identical laser fluence has been observed. Since the dynamics of the polymerisation are supposed to be the reason for this, further studies are planned to this issue.The curing depth cd for Ormocer b59 with 0.5% of ADM 3 and varying parts of MAM (0%-50% has been measured at a hatch distance of dh = 0.001 mm. The result is depicted within figure 11. The graph illustrates that the curing depth slightly increases with increasing part of MAM.

The highest process resolution was achieved without MAM (18 μm) in contrast to 20 μm when using 50 % of MAM at a laser fluence of H0 = 20 J/cm².1910080m μ /dc h t p e d g n i r u c604020Ormocer b59 ADM 3 0,5% Ormocer b59 ADM 3 0,5% MAM 10% Ormocer b59 ADM 3 0,5% MAM 50%01x1032x1033x1034x103laser fluence H 0/ J cm-2Figure 11:The curing depth cd versus the laser fluence H0 of modified Ormocer b59 ADM3 0.5% with varying parts of MAM (0-50%) at a hatch distance of dh = 0.001It has been further found, that the chemical cleaning of the produced structures is more effective, the more parts of MAM are introduced in the basic material.Investigations on three dimensional structuring:Increased complexityFigure 12:Investigation strategy during the fabrication of micro mechanical systems: CAD models of basic modules with high aspect ratio towards movable micro mechanical systemsThe investigations which are in the following presented have been made with Ormocer b59 and 0.5% of ADM3 at a viscosity of 3.2 MPa.

(0% MAM). The main challenge during the investigations was the transfer of the process parameters which have been found during 2D and 2 ½ D test towards the 3D patterning of micro mechanical systems and, if necessary, their additional fine adjustment. In order to apprehend systematically the parameter effects on the process, CAD models with varying level of complexity (figure 12) have been developed. In a first instance, basic modules have been planned to demonstrate the capability of the selected material and the custom made machining setup for the layer-based production of three-dimensional structures. In the following, basic investigations for the direct production of micro mechanical20systems having moving parts without the necessity of assembly steps have been done.

In the following, basic investigations for the direct production of micro mechanical20systems having moving parts without the necessity of assembly steps have been done. Based on these results, the fabrication of functional parts with increased complexity for micro mechanical applications will be studied.Results During the investigations on three dimensional structuring, it has been shown, that the results of the curing measurements are only conditionally useful. In fact, for achieving a high contour accuracy, a careful adjustment of the fabrication parameters has to be done, mainly depending on the complexity and design of the structure to be built. In the first instance, basic cylindrical structures with a high aspect ratio up to 10 (Figure 13 (a)) have been successfully fabricated to show both the machining capability and the suitability of the developed material for micro structuring.

In order to ensure a reliable adhesion between individual H0 > 400 J/cm² at a hatch distance of dh = 0.01 μm.layers,the first experiments have been carried out with alaser fluencebFigure 13:High aspect ratio (HAR) structure developed with the custom made stereo lithography setupThe structure displayed in figure 13 (b) has been fabricated with a laser fluence of H0 = 560 J/cm² and consists of 250 individual layers at a thickness of 10 μm. The outer diameter of the pillar was 400 μm. In the following, the direct production of complex structures with movable parts has been investigated. For this, the CAD model depicted within figure 14 (a) has been developed. Basic approach was the fabrication of one basic structure that adheres at the building substrate. The components that will be movable after the finished fabrication will be carried by cylindrical support structures to prevent their displacement during the production cycle.

The components that will be movable after the finished fabrication will be carried by cylindrical support structures to prevent their displacement during the production cycle. Subsequently, the dissolution of the support structures is planned to be done automatically during the chemical cleaning process since other removal techniques are not reasonable in the micro meter scale.abcFigure 14: aAchieved results when fabricating micro parts with H0 = 560 mJ/cm², dh= 0.01 mm (b), and H0 = 4 x 103 J/cm², dh= 0.001 mm (c)21With varying process parameters (laser fluence H0, hatch distance dh) the optimum fabrication parameters have been distinguished. It can be generally mentioned, that decreased laser fluence H0 enhances the process resolution, but minimizes the surface quality at a constant hatch distance dh. Figure 14 (b) shows clear lateral gaps, but the surface has a periodic structure.

Figure 14 (b) shows clear lateral gaps, but the surface has a periodic structure. The structure depicted in figure 14 (c) shows only minor contour accuracy, but simultaneously high surface quality. Further process optimisation lead to the result depicted within figure 15.Figure 15:Achieved results when fabricating micro parts with H0 = 500 J/cm², dh= 0.001 mmClean gaps and very smooth surfaces are visible when using H0 = 500 J/cm² and dh= 0.001 mm. The support structures are well-defined with a diameter < 40 μm. The further challenge was the separation of both ring and basic structure as key assumption for the production of movable parts. It has been shown, that the stability of 10 μm supports for carrying ring-structures was sufficient during the process to withstand the mechanical stresses of the coating procedure on the structure. Preliminary tests of dissolving the support structures have shown success.

Preliminary tests of dissolving the support structures have shown success. The ring-structure has been individually separated from the basic structure (figure 16)Figure 16:SEM and microscope images of two separated structures that have been fabricated in direct production without any assembly steps. The support structures are chemically dissolved during the cleaning process.In the following the focus was on the production of structures for micro mechanical applications. Micro systems with evolvent gearwheels have been produced (Figure 17). The fabrication parameters have been using H0 = 400 J/cm², and dh= 0.001 mm. The gearwheels have been conjoined with axes that were on the other hand only conjoined with the bottom of the basic structure by supporting cylinders. The axes were specially embedded at their bottom and top inside precursors of bearings. Inside the base of the structure, drainage openings have been constructed to allow the chemical cleaning of the lower bearing.

Inside the base of the structure, drainage openings have been constructed to allow the chemical cleaning of the lower bearing. It has been shown, that the transmission of rotations is possible.22Especially the meshing of the gearwheels has been fabricated with distinctive quality. The distance between individual teeth is in the order below 10 μm. The fabrication of basic drive units has been investigated to drive the micro systems with an air or water flow for micro fluidic applications.Figure: 17Micro mechanical systems with technical gearing systemFigure 18 shows the first results when fabricating water wheels. Water shovels with a lateral size of 680 μm have been produced. Lateral gap distances between axle and bearing were successfully built below 50 μm.

Water shovels with a lateral size of 680 μm have been produced. Lateral gap distances between axle and bearing were successfully built below 50 μm. Drive tests have been successfully accomplished with a slightly modified structure that has been built up symmetrically by arranging the wheel between lower and upper bearing.Figure 18:Functional micro mechanical assembly with water wheel for fluidic applicationsConclusion and outlook Within this paper, the authors have presented the development of a custom made stereo lithography setup. In combination with new materials suitable for rapid manufacturing, functional micro parts have been built up. By using the classical Rapid Prototyping approach, parts have been directly built in a layer based production process. Polymer micro systems have been fabricated with both lateral and vertical resolutions below 10 μm.

Polymer micro systems have been fabricated with both lateral and vertical resolutions below 10 μm. The developed One Step Production technique (OSP) enables the production of micro mechanical assemblies with included moving parts. Here, the in-process assembly without the necessity of further conjoining steps of the functional parts is one of the main innovations. Besides resolutions in the micron range with high surface qualities, one key approach towards Rapid Manufacturing is the use of functional polymer materials with mechanical properties suitable for functional parts.23During the fabrication of three dimensional structures, the high-viscosity resin has been successfully applied. However, a long settling time has lead to long process durations. We have shown the possibility of the chemical modification of the basic material to reduce effectively the viscosity.

neration of Porous Structures Using. Fused.D~positionBertoldi M., Yardimci M. A., Pistor C. M., GUyeri S. 1., Danforth S. C. University of Illinois at Chicago & Rutgers UniversityABSTRACTThe Fused Deposition Modeling process uses hardware and software machine-level language that are very similar to that of a pen-plotter. Consequently, the· use of patterns with poly-lines as basic geometric features, instead of the current method based on filled polygons (monolithic models), can increase its efficiency.In the current study, various toolpath planning methods have been developed to fabricate porous structures. Computational domain decomposition methods can be applied to the physical or to slice-level domains to generate structured and unstructured grids. Also, textures can be created using periodic tiling of the layer with unit cells (squares, honeycombs, etc).

Also, textures can be created using periodic tiling of the layer with unit cells (squares, honeycombs, etc). Methods 'based on curves include fractal space filling curves and.change of effective road width Within a layer or within a continuous curve. Individual phases can also be placed in binary compositions.In present investigation, a custom software has been developed and implemented to generate build files (SML) and slice files (SSL) for the above-mentioned structures, demonstrat ing the efficient control of the size, shape, and distribution of porosity.INTRODUCTIONTwo dimensional pattern generation capability of RP/SFF .methods is not being fully utilized currently due to use of very .limited set of toolpath generation algorithms and data representation problems associated with heterogeneous structures.

Generation .ofcontrolled porosity structures with RP/SFF has been particularly challenging due to the need of defining the entire porous structure as a very complicated solid object.Porosity generation within solid models with RPISFF .• processes has been investigated in the context of investment casting with ·•SLA •fabricatediconstantpcrtesize patterns. ••• Two" dimensional equilateral· triangular and square self-repeating patterns were developed by Systems, three-dimensional laminated hexagonal structures>by the University of Nottingham group [1] and three-dimensional tetragonal crystalline structures by Gervasi and coworkers [2].639Recently, rectangular porous interior structures were adopted by Stratasys in Fused Deposition and tetragonal patterns by Z-Corporation in 3D printing for build acceleration. The comp~tational geoIIletry •.

The comp~tational geoIIletry •. definition problem of heterogeneous.structures has been investigated [3]; however, this study does not address the generation of porous structures inside solid models.:Fused Deposition can be consid~red as a vector printing technique, in which the two dimensional patterns are generated with poly-lines on a discretized x-y plane. Boundary deposition and subsequent raster filling is employed during production; voids appear in the raster filling, due to the finite road width; .this can be used to produce specific porosity. However, current Fused Deposition process plannin.g software only allows for constant layer thickness across the entire part, constant road width within each layer, contour or rectilinear rasters, monolithic material composition, and constant air-gap between raster patterns. As a result, the variety of attainable porous structures is very limited.A set of software tools has been.

As a result, the variety of attainable porous structures is very limited.A set of software tools has been. developed and used to generate porous structures, showing the feasibility of the concepts presented. Since layer geometries are inherently two the remainder of the presentation will be focused on two-dimensional grid dimensional, generation.TWO DIMENSIONAL TEXTURESBasic texture structures employed by computer graphics methods, such as square blocks and honeycombs, can fill a plane by tiling the basic unit periodically. The variety of micro structures can be increased by recursively constructing higher-order unit cells that fill the plane through periodic tiling, as·shown in Figure 1.oFigure 1 Unit cell and resulting texture.is designed properly to fill the entire plane. The texture needs to be clipped to fit into the layer geometry described by the contours, which increases the computational requirements.

The texture needs to be clipped to fit into the layer geometry described by the contours, which increases the computational requirements. The unit cell nature of the textures enforces a homogeneous pore size and microstructure distributions throughout the entire part, which limits the use of this technique.Texture based patterns can be generated once the underlying unit cell640STRUCTURED GRID GENERATIONStructured grids are used in the simulation of complex physical processes, where discretization of the domain of interest into geometrically simple sub-structures is required. Systems of algebraic or ordinary differential equations representing the physical phenomena are defined over the discretized domain, with appropriate variable transformations [4].

Systems of algebraic or ordinary differential equations representing the physical phenomena are defined over the discretized domain, with appropriate variable transformations [4]. Uniform Cartesian grids within rectangles and circular geometries can be generated, employing linear functions and coordinate transformations, as depicted in Figure 2..1000~~IOOO500o500o500Figure 2 Uniform Cartesian grids.The resultant grids have regular porosity distributions, which can be manipulated only by varying grid spacing independently along the coordinate axes. The porosity can be better controlled through use of nonlinear stretching functions as mapping functions from computa tional to physical space.

Complex geometry physical simulation641(1)necessitated the development of boundary conforming, or boundary fitted curvilinear grid methods [5]., . . - - - - - - - - - , 300020001000 oy10002000xFigure 3 Non-uniform Cartesian grids with boundary clustering.Boundary conforming structured grids can be efficiently produced by using algebraic interpolation methods. Portions of domain boundaries in physical space are associated with domain edges in computational space. Interpolation methods are utilized to generate coordinates of interior grid nodes; a two-boundary interpolation of a circle is presented in Figure 4.3000250020001500 Y1000500Figure 4. Boundary conforming grid generated with two-boundary interpolation.Greater flexibility of boundary conformity can be obtained through higher order interpolation algorithms, e.g. transfinite interpolation:r.

transfinite interpolation:r . = pi + pj _ pij ',Jpi =h f;m,j +(1- h)~,j, pj = gj f;,jm +(1- gj )f;,lpij = h gj f;m,jm +h (1- gj )f;m,l +(1- h)gJ"jm +(1- h )(1- gj )~,l}-1 Ii =-.-1 ' gj =-'-1 ;m-i-I zm-(2)The resultant grid for the same circular geometry is depicted in.Figure 5. Regions around which grid nodes are concentrated and cells are distorted, correspond to the comers of the rectangle in physical space.Porosity control on algebraic.. grids can be further enhanced through non-linear stretching functions or advanced mapping methods from the theory of complex variables. As an example,.a simply connected .arbitrary geometry polygon can be mapped onto a unit circle in the complex plane, i.e. computational space, through a series of conformal maps [6],[ 7].Generation of boundary conforming curvilinear grids can also be formulated as a boundary or initial/boundary value problem (IVBP).

computational space, through a series of conformal maps [6],[ 7].Generation of boundary conforming curvilinear grids can also be formulated as a boundary or initial/boundary value problem (IVBP). IVBP's can be described by partial differential equation (PDE)as the governing equation over the domain and relevant initial and boundary conditions, where dependent variables are coordinates of the grid points.

Since a variety of different PDE's· can be used, greater control is achieved on pore sizes, clustering, continuity and orthogonalitypf grid lines.3000250020001500 Y1000500Figure 5: Boundary conforming grid generated with transfinite interpolation.Mapping between computational and physical spaces is defined with a variable transformation, for effective construction of PDE based grid .generation methods:x =X(4'",17) 4'"= 4{'x,y), Y=Y(4'",'l) , 17= r,{x,y){:~} =(x,y, ~x,y,l:;, -x:'H:}where subscript denote partial differentiation of the dependent variable with·respect totheinde pendent variable.

The most widely used 'PDE's are second order elliptic equations, written in the canonical form:643(3)For grid generation, a system of elliptic PDE' s can be constructed with grid coordinatesas dependent variables in respective problem spaces:c1- ~+ c1- ~ = p oY ox22c1- '7 oxc1- '7 Q oY, -2 +-2 =,(ti-) J-)) _(;:) .",'7 = S\r ,,/\rr Erporwhere (P,Q) are grid control parameters in physical space with their counterparts (<I>,<p) in computational space. Since Dirichlet (constant value) type boundary conditions are specified for the BVP,the elliptic PDE'sonly needto be solved in the interior of the domain.

6), was set to 1.8 and iterations were stopped when the average cell-based residual was calculated to be less than % 0.1. The resulting grids arepresented in Figure 6.- - - - - - , 4000 3500 3000 2500 2000 1500 1000 500Y~---__."..,"T'I' . . , . . . - - - - - - , 4000 3500 3000 2500 2000 1500 1000 500Figure 6 Curvilinear grids generated with second order elliptic PDE's: no clustering (a), clustering (<I> = 0.03) in ~ direction (b).MULTI-GRID METHODSIf the solution of a computational problem requires higher grid node densities at specific locations within the domain, it is possible to generate an additional fine grid. Embedded grids, as shown in Figure 7-a, provide local density variations without distorting the base grid; however, the complexity of the algorithms is increased due to additional interpolation/interrogation steps.

In analogy to computational physics simulations, embedded grids may provide refined porous structures.2000150010005004000 3500 3000 2500 2000 y 1500 1000 500050010001500200001000300tPFigure 7 Sample embedded grid (a) and multi-block grid (b).645(8)For complex •.. geometry domains mapping of the Cartesian grid may result in severe distortion and low quality curvilinear grids, regardless of the particular method employed. Multi a4dress grid quality bygec0tnposingthe domain into simpler geometry sub block.methods domains, .generating grids •within each of them using any of the techniques described above. Multi-block grids are hybrid geometrical entities; each sub-domain is made of a structured grid but the overall grid topology is described in the context of unstructured grids (Figure 7-b).

Multi-block grids are hybrid geometrical entities; each sub-domain is made of a structured grid but the overall grid topology is described in the context of unstructured grids (Figure 7-b). Multi-block grids offer increased geometric flexibility over single block structures; however, the algorithms are more complex and mostly semi-automatic.iiUNSTRUCTURED GRIDSUnstructured grids discretize the problem domain into basic geometric features that do not have implicit topology. Tl1eir connectivity is defined explicitly as •part of the data-structure, in .contrast to structured grids. As ares~lt, computational lllet119ds are less efficient due togenerationmethogs are geometrically more flexible and robust, since the specification of domain boundaries with closed contours (2D), or closed surfaces (3D), is· sufficient for initial grjd generation [8]. Triangular grids are the most widely used type of unstructured grids and can be generated with different methods:..

Triangular grids are the most widely used type of unstructured grids and can be generated with different methods:.. Mapping of a pre-triangulated mesh onto the domain using structured grid generati()n methods, or conversion of two-dimensional structured grids, made of quadrilaterals, into triangles... Delaunay triangulation, which generates triangles for a pre-specified grid •. node distribution optimally. Optimality is defined with the conformity of the generated triangles to equilateral triangles [9]... Advancing front method, in which the domain boundaries define a starting front. In contrast to Delaunay triangulation, advancing front method generates interior point cloud using a set of heuristic rules [10].

In contrast to Delaunay triangulation, advancing front method generates interior point cloud using a set of heuristic rules [10]. Parameters and transformations within the rules can be manipulated to obtain construct tailored triangulations with relatively high computational efficiency [8].A dual tessellation of the>polygon interior can be obtained for unstructured grids by generating new points at triangle center-points (type I), or by observing distance constraints (Voronoi tessellations, type II) and constructing polygons around the old grid nodes. Interior points of Voronoi polygon Vi are closer to old grid node Pi than any other node, Pj, in the original unstructured grid. For ideal equilateral triangular grids, both constructions produce identical hexagonal dual tessellations. For· arbitrary .grids there· are considerable differences,646where type I reflects the underlying triangulations signifi<;antly better. Figure 8 depicts.

For· arbitrary .grids there· are considerable differences,646where type I reflects the underlying triangulations signifi<;antly better. Figure 8 depicts. a constrained Delaunay triangulation,a dual center-point tessellation and the relative layer..wise superposition for a triply connected domain, showing as these (I.lgorithnls can be effectivc::lyutilized for porosity generation.25002500250020002COJ20001&101&101500100010001000500500&10025000&1010001&1020002500Figure 8 Triangulation (a), tessellation (b) and superposition of the two (c) on a triply connected domain.DOMAIN DECOMPOSITIONThe capabilities of individual porosity generation methods can be further amplified utilizing domain decomposition tools on layer level geometry contours. Distributions of raster angles, road widths and air-gaps can be assigned to each sub-domain by automatically modifying the properties of individual polygons in the slice (SSL) file [9].

Distributions of raster angles, road widths and air-gaps can be assigned to each sub-domain by automatically modifying the properties of individual polygons in the slice (SSL) file [9]. Although domain decomposition is not as geometrically flexible as direct toolpath generation techniques described above, it can be integrated into existing process planning chain more easily.SPACE FILLING CURVESSFC's can be utilized as novel toolpath generation tech~iques inthe generation~t~orous structures. The lengthoftheunit steps within SFC's can be settoavaluelargerthanWidthofthe roads, to generate porosity. Three..dimen.si()nal porosity can be fabricated by st(l.ckingtotatedor mirrored patterns of the origin'll layer with respect to symmetry axes. See reference [9]. for a more detailed description and someex.amples.MULTI-MATER.lAL DEPOSITIONFuse.d Deposition hardware (e.g.

See reference [9]. for a more detailed description and someex.amples.MULTI-MATER.lAL DEPOSITIONFuse.d Deposition hardware (e.g. StratasysFDM 1650) is able to deposit two different materials (modeling and support), each of which has separate delivery systems and flow6472500controllers. In SML language, the material for each poly-line can be specified. Hence, it is possible to specify material distributions for the porous structures inside the toolpath generation algorithms.

Hence, it is possible to specify material distributions for the porous structures inside the toolpath generation algorithms. This can be done in different ways:Randomly, either by assigning the material to a particular polygon through random number generators, or using center-point coordinate of the polygon and utilizing two-dimensional binary random patterns.With a real valued function on a plane,f(x,y): the function is evaluated at the center-point of individual polygons and the material is selected based on the digitization of the calculated function.Figure 9 shows a part that has been fabricated based on an unstructured grid of a simply connected domain and having a material distribution assigned according to a function specified asf = cos(_xJSin(~J10002000Figure 9 Multimaterial porous structure.GENERATION OF POROUS STRUCTURES WIH FUSED DEPOSITIONPorous structures are described with open or closed poly-lines.

A routine has been developed for the translation of poly-lines with appropriate properties into SML files [9]. In contrast to transformation based RP/SFF methods, e.g. SLA andSLS, PD necessitates the use of non-intersecting toolpaths. The minimum feature size in FD is defined by nozzle diameter and attainable road width range is currently 0.014"-0.040" for the standard 0.012" nozzles. Since one micro-step corresponds to 0.001" in x-y plane, the smallest controllable deposition is approxi mately 0.130" long. Although these bounds can be lowered using low deposition head speeds,648the minimum deposition length will not be shortened below 0.1" in the existing material delivery system, due to considerable inertia of the melt within the liquefier.Also, due to limited communication between motion and flow controllers, straight lines, poly-lines with comers, and corners with small and large angles are not differentiated.

As a result, excess material will be generated invariably at internal side of the comer, possibly closing the>pores.For structures that rely on deposition of unsupported bridges, maximum pore size has been estimated as 0.150" for ABS material. However, for self-supporting structures the.pore size is only bounded by the size of build domain.SCAN-LINE DEPOSITIONStructured grids and space filling curve patterns can be effectively achieved through deposition of open-ended poly-lines. Although structured grids can also be generated using polygon deposition methods, individual grid lines within the structures are natural building blocks for scan-line deposition. Structured grids can be generated via a number of ways:Families of grid •.

Structured grids can be generated via a number of ways:Families of grid •. lines, ~ and 11, can be deposed on separate layers resulting in openporous structures along the side-walls, or on the same layer, resulting in laterally closed cells and excess material accumulation at grid nodes.lines are Staggered grids can be inserted into layer stacking type a, produced on alternating layers. Considerably reduced sagging is expected to occur in comparison to type a, due to reduced bridging lengths.in which ~ and 11Staggered grid ~ and 11 lines are produced on the same layer. Although the overall structure will have open porosity along the wall surfaces, each layer will have closeclboundaries.CONCLUSIONSThe work has presented the possibility of using Rapid Prototyping for the production of porous structures, using algorithms based on different computational approaches.

Software tools have been developed and a number of porous structures has been built using a StratasysFDM 1650 machine, shoWing the feasibility of all the proposed concepts. The future work wilFinclude use· of polygon renumbering methods which. are required for polygon-deposition of unstructured grids or their~ualtessellationtoshortenthedeposition time as well as field adaptive unstruCtur ed.gridgeneration techniques.649REFERENCES1.Hague R., Dickens P., "Design of new build structures for the successful autoclaving of stereolithographymodels," The Seventh International Conference on Rapid Prototyping, pp. 192-202, March 31-April 3 1997,San Francisco, California, USA.2.Gervasi V. R., BrandfD.A., Shaffer S.D.,Lim K., "Tetracast SLAbuHclstyle," The Seventh InternationalConference on RapidPrototyping, pp.309-317, March 31-Apri131997,.SanFran6isco, California,·USA.3.Sun W., Lau A.

nerative Design for Additive Manufacturing of Satellite Optical Tracker MountPaula Logozzo1, Donald Palomino1, Abraham Meiszner1, Bodia Borijin1,2, Andrew Wang1,3, Ryan Watkins4, Bingbing Li1,51Autonomy Research Center for STEAHM (ARCS), California State University, Northridge, CA 91324 2Department of Structural Engineering, University of California San Diego, La Jolla, CA 92093 3Portola High School, Irvine, CA 92618 4Jet Propulsion Laboratory, California Institute of Technology, Pasadena, CA 91109 5Department of Manufacturing Systems Engineering and Management, California State University, Northridge, CA 91330AbstractThe organic and intricate nature of machine generated parts is problematic during the manufacturing phase, resulting in high costs and slow production with traditional manufacturing techniques. Additive manufacturing has been explored and approved as the potential solution for fabricating complex geometry and organic lattice structures.

Additive manufacturing has been explored and approved as the potential solution for fabricating complex geometry and organic lattice structures. With the advances in topology optimization and generative design, design for additive manufacturing (DAM) allows users to generate numerous, high-quality design alternatives that are lighter and stronger than traditionally designed parts. This study addresses the process of designing and load testing a satellite optical instrument mounting bracket using the generative design and simulation capabilities of Autodesk Fusion 360.

This study addresses the process of designing and load testing a satellite optical instrument mounting bracket using the generative design and simulation capabilities of Autodesk Fusion 360. The workflow pipeline begins with the creation of generative design studies for the instrument bracket in accordance with the design criteria outlined in the optical instrument design challenge, where it is then thermally load tested using finite element analysis (FEA) in Fusion 360 and analyzed for its mechanical behavior.Keywords: Generative Design, Design for Additive Manufacturing, Satellite Optical Tracker MountIntroductionWith extreme temperature fluctuations typically spanning hundreds of degrees Celsius, the environment of space is one of the harshest places for machinery to operate.

Satellites orbiting the earth will undergo frequent wide temperature swings from -140℃ to +140℃ when transitioning between the earth’s shadow and unfiltered sunlight, resulting in a wide operational temperature range for all components on board [1]. This wide operational temperature range has a dramatic influence on the geometry of every solid component on the satellite due to thermal expansion. Geometric changes are problematic for a variety of reasons; they can cause irregularities in the performance of moving parts, and they can also interfere with the accuracy of finely tuned onboard optics and communications instruments. Thermoelastic geometry changes are unavoidable to a certain extent - all materials have a coefficient of thermal expansion (CTE) that, for a material exposed to a thermal load, will result in some amount of geometric deformation.

Previously, engineering teams responsible for the design and manufacture of satellite components attempted to address this problem through improving the thermal properties of component materials in order to reduce thermoelastic deformation as much as possible.2081Literature review1) Prior satellite component manufacturing methodsA ceramic composite material, HB-Cesic, was successfully used by a team of engineers in the design and manufacture of a star tracker assembly bracket on board the Meteosat Third Generation (MTG) weather satellite, operated by the ESA [2]. HB-Cesic was chosen because of its optimal thermoelastic properties, especially in comparison to materials such as Aluminum 6061, Invar, and Zerodur. With a density of 2.9 g/cm3, a thermal conductivity of 200 W/mK, and a CTE of 2.3 e-6/K, HB-Cesic is an attractive material for the construction of satellite components [3].

With a density of 2.9 g/cm3, a thermal conductivity of 200 W/mK, and a CTE of 2.3 e-6/K, HB-Cesic is an attractive material for the construction of satellite components [3]. Satellite optical system components have also been successfully constructed out of another thermally optimal material, Zerodur, and used on the FOKUS, KALEXUS, and MAIUS-1 rocket missions [4]. Benefits of using these materials include improved thermo-mechanical stability and a reduction in thermoelastic deformation under thermal loads [2, 4]. However, one of the major drawbacks with the usage of components constructed from HB-Cesic or Zerodur is that it relies on traditional manufacturing methods such as CNC machining, trimming, and assembling [2, 4]. Additionally, current design methods are manual in nature and do not use automation in the design process.2) Automation for designAutomation as a design tool is a relatively new technology in the field of engineering.

Traditional design methods typically require conscious and deliberate input from a person, or team, at every step of the design process. While many engineering endeavors still require constant human oversight, certain elements of the design process can now be automated for the first time. Automation allows engineers to explore novel solution possibilities for a great variety of engineering problems while also frequently saving time and potential costs during the design phase. All too often, engineers can become entrenched in particular solution methods that limit their innovative and creative abilities; this has the effect of limiting the conceptual range of possible design solutions for an engineering problem. Automation tools such as generative design and topology optimization can help overcome this issue by providing a wide range of novel solutions for engineering problems that surpass the limits of human imagination.

In brief, and with respect to design for manufacturing (DFM), topology optimization can be characterized as a kind of iterative mass subtraction for the purposes of weight reduction and removing unnecessary mass, whereas generative design can be characterized as the iterative rearrangement of mass, and potentially the addition of mass, to optimize a part or component with respect to its design requirements and boundary conditions.Generative design is quickly being adopted by research teams as a powerful automation tool with a wide range of usage cases. It has been successfully used for the layout and planning of a Swedish residential city block that is optimized for space efficiency and minimizes the environmental impact of construction [5]. It was also used alongside Generative Adversarial Networks (GAN) for image-based design of garments for the fashion industry [6].

It was also used alongside Generative Adversarial Networks (GAN) for image-based design of garments for the fashion industry [6]. Other usage cases include the creation of a series of 6,010 virtual buildings with randomized layouts for the purposes of analyzing the effect of structural geometry on the buildings’ energy consumption [7]. There is a growing body of evidence demonstrating that these automation solutions have a broad usage range and a tremendous potential to be used alongside AI-based computing software for mechanical and structural designs. In a recent case study to evaluate the effectiveness of a design methodology for modular spacefaring equipment, an engineering team used topology optimization2082alongside genetic algorithms to design a satellite antenna [8]. The team was able to construct an algorithm that factored in a number of design and manufacturing criteria including manufacturing cost, design adaptability, and post-processing cost, among others.

The team was able to construct an algorithm that factored in a number of design and manufacturing criteria including manufacturing cost, design adaptability, and post-processing cost, among others. These criteria, paired with topology optimization, produced a range of unique antenna architectures in which the design of the antenna was optimized for each criterion [8]. Another research team successfully integrated topology optimization into a generative design framework to quickly produce a wide range of structural design outcomes that can be evaluated according to different parameters such as cost, novelty, and compliance [6].

This design methodology pairs elegantly with additive manufacturing techniques, which are quickly becoming one of the dominant forms of manufacturing for space applications [9].3) Simulation capabilities with generative designAdditive manufacturing is an attractive option for the manufacture of spacefaring components because of its mass-reduction capabilities, as well as its cost-saving potential due to the large amount of material waste associated with traditional subtractive manufacturing techniques [10,11]. It is estimated that additive manufacturing can reduce the cost of a titanium alloy component by as much as 50% due to the reduction in material waste from subtractive manufacturing and machining techniques [12]. Additionally, the ‘Buy-to-Fly’ ratio of spacefaring components, which is the ratio of purchased material stock mass to final component mass, is typically anywhere from 10:1 to 25:1, and can even be as high as 40:1 [13].

3D metal printing has been shown to bring this ratio down to 1.5:1, and even as low as 1:1 due to the net shape component geometry produced by additive metal manufacturing [14]. Compared with traditional subtractive manufacturing, additive manufacturing is often faster; this reduction in lead time can reduce the costs of manufacturing satellite components [15]. Additive manufacturing techniques also allow for the mass production of highly intricate, geometrically complex mechanical structures that are difficult to create using subtractive manufacturing [16]. Bugatti was recently able to produce a powerful new brake model for the Chiron by employing selective laser melting (SLM) of a Ti- 6AL-4V powder; this brake model was shown to be stronger than prior models with a 40% reduction in mass compared to previous designs [17].

Current industry trends suggest that SLM, as a manufacturing technique, is quickly becoming the dominant component additive manufacturing technique in the aerospace industry; this makes SLM the most likely technique for the manufacture of a satellite mounting bracket like the ones generated in this study [18]. SLM was used to fabricate an aircraft component that was redesigned using topology optimization, with an overall volume reduction of 54% and a weight reduction of 28% between the original aluminum part and final titanium alloy optimized part [19]. Since the geometries produced by generative design for topology optimization are often intricate and organic in nature, it follows that the most efficient way to produce generative design models is through additive manufacturing.Overview of ProjectPurposeOptical systems are some of the most commonly launched instruments in the space industry.

Some prominent examples of optical instruments in space include the Hubble Telescope and the newly launched James Webb Space Telescope, however optical instruments can also be found on Mars rovers and even aboard small CubeSats in low earth orbit. In order for these instruments to capture deep regions of space with any degree of accuracy, optical systems have2083very strict pointing requirements. Any deviation in the pointing direction of the tracker lens will result in tremendous photographic inaccuracies for deep space photography applications.Outline of design requirementsThe pointing accuracy of an optical system is defined by the ‘centerline’ of the lens, which can be visualized as a straight line pointing through the central axis of the lens. Since optical instruments are typically mounted on a rigid surface, it is expected that the centerline will not move or shift over time.

Since optical instruments are typically mounted on a rigid surface, it is expected that the centerline will not move or shift over time. However, optical instruments on satellites orbiting the earth are subjected to intense thermal loads due to sunlight exposure; this creates large temperature differences in the mounting material between unfiltered sunlight and the full shadow of the earth. As a result, the mounting material will expand and contract due to thermal expansion. This expansion and contraction results in thermoelastic deviation of the central axis of the lens, which reduces the accuracy of the optical system over long distances. It is therefore crucial to design a mounting bracket that expands and contracts in such a way as to minimize the angular deviation of the centerline.Outline of design constraintsDesigning a mounting bracket that deviates as little as possible under thermal load is a current problem faced by engineers in the space industry.

This design problem, posed by Dr. Ryan Watkins of JPL, challenges engineers with the task of designing an optical tracker mount that deviates minimally while subjected to thermal loads. The main goal for the tracker design challenge is to produce a mount with a centerline that deviates no more than 0.001° away from its original position. The mount must also have a minimum cross-sectional member size of 1 mm, a fundamental frequency greater than 200 Hz, and a minimum factor of safety of 1.25. The design challenge poses several loading scenarios for the tracker mount, including a force of 2,000 N in the x, y, and z directions, a bulk temperature soak of 85℃, and a steady state heat conduction condition with the tracker at 50℃ and the base pads at 10℃. Additionally, all mounting pads along the base are treated as fixed in space. A design space envelope is given as a volumetric reference; the final design must fit within the dimensions of the envelope.Methodology1.

A design space envelope is given as a volumetric reference; the final design must fit within the dimensions of the envelope.Methodology1. WorkflowTo meet these design challenges, our team utilized the generative design capabilities of Autodesk Fusion 360 to construct the mounting bracket. Currently, there are very few commercially available automated design tools that are both effective and user-friendly. Fusion 360 aims to meet this need by providing powerful automation design tools with an intuitive user interface. The goal of this research is to use Fusion 360 generative design to produce a bracket design that thermoelastically deforms in a controlled manner. Previously discussed satellite components, such as those constructed from exotic materials such as HB-Cesic or Zerodur, attempt to address the problem of thermoelastic deformation through the net reduction of deformation by using materials with low CTE values.

This research attempts to address the problem of thermoelastic deformation by using automation to create a series of titanium satellite bracket designs that deform in a predictable manner to reduce the angular pointing deviation of a satellite optical system. This method employs comparatively cheaper materials, automation in the design process, and additive manufacturing. These methods combined can potentially reduce material2084cost, material waste, and production lead time, resulting in globally reduced costs for functionally equivalent satellite components.The workflow for this project is organized into two sequential steps: generative design and thermal load testing. The generative design phase begins with the design space for the model, which is defined by the dimensions shown in Figure 1. The bracket model created during the generative design phase must fit within the volume envelope defined by the design space dimensions.

The bracket model created during the generative design phase must fit within the volume envelope defined by the design space dimensions. Once the model is completed, it is exported to a 3D format for thermoelastic deformation testing. The model is subjected to a series of thermal loading scenarios, and subsequently analyzed using FEA for its behavior under each loading condition. Models that do not meet the design requirements for thermoelastic deviation are rejected as potential model candidates.Fig. 1: a) Design space dimensions for the optical tracker mount, b) Tracker CAD modelDuring analysis, it is imperative to keep a static reference point for the optical trackercamera itself. This is shown in Figure 1 b, which is a model for the optical tracking instrument.

This is shown in Figure 1 b, which is a model for the optical tracking instrument. The tracker model is used as a geometric and spatial reference point during thermal FEA, and also serves as a fixture point for the upper mounting pads.Figure 2 shows the design space as it appears when the tracker model is fixed to the front. Note the instrument mounting pads, shown in green, beneath the optical instrument model. These mounting pads are not valid design spaces because they are an off-the-shelf component and therefore cannot be altered. The yellow pads on the bottom of the design space, however, are free to be altered to produce novel bracket designs. Altering the base mounting pads is one of the first crucial steps in generative design input, which is outlined below.2085Fig. 2: Optical tracker model placed on design space, shown in violet2. Generative designInputs for Fusion 360Fig.

2: Optical tracker model placed on design space, shown in violet2. Generative designInputs for Fusion 360Fig. 3: a) Design space with unassigned bodies, b) Preserve geometry assigned to bodies, c) Obstacle geometry assigned to shell, d) Section view of shell in design spaceThe first step in utilizing generative design is setting up a design space consisting of bodies to later be defined as preserve and obstacle geometries. An example of this is shown in Figure 3a)2086where the gray bodies were created using the Design workspace within Fusion 360. Unlike conventional part design, generative design only requires bodies that will be used as preserve or obstacle geometrics. Therefore, a body like that shown in Figure 1 is not required to be made. Not shown here is the shell body that will later serve as the obstacle geometry.

Therefore, a body like that shown in Figure 1 is not required to be made. Not shown here is the shell body that will later serve as the obstacle geometry. Once all bodies are created within the Design workspace, the rest of the setup process will continue in the Generative Design workspace.Multiple inputs and parameters must be defined in the Generative Design workspace before a study can be successfully conducted. First is assigning the bodies made in the Design workspace as preserve or obstacle geometries. As their name implies, preserve geometries are assigned to bodies to incorporate them in the final shape of the design, while obstacle geometries are assigned to bodies to represent spaces for the design to avoid. These geometries are assigned within the Generative Design workspace once the bodies in the Design workspace are finished. Figure 3b) shows preserve geometries assigned to the bodies representing the optical tracker and eight mounting pads.

Figure 3b) shows preserve geometries assigned to the bodies representing the optical tracker and eight mounting pads. Figure 3c) shows the obstacle geometry assigned to the shell body surrounding the optical tracker and mounting pads. A section view of the interior space inside the shell is shown in Figure 3d). This represents the design envelope defined in the problem statement. This ensures that the generative design results do not exceed this defined envelope. Lastly, a symmetrical plane is created through the design space in the XZ plane to produce symmetrical generative design results. This lateral symmetry is beneficial for the thermal analysis and will be further elaborated upon in a later section.Once all of the preserve and obstacle geometries have been assigned, design conditions including structural constraints and loads are defined. Structural constraints are to apply fixes, pins, or frictionless constraints to a geometry.

Structural constraints are to apply fixes, pins, or frictionless constraints to a geometry. In Figure 3b), four lock icons can be seen on the bottom four mounting pads, indicating a fixed structural constraint. Structural loads simulate pushing, pulling, and twisting forces. The force identified from the problem statement is applied onto the optical tracker by defining it as X, Y, and Z components of a vector. It is also important to remove the gravity force that is included by default.Design criteria, such as design objectives, manufacturing method, and materials are also defined. Design objectives allow for the generative design study to prioritize minimizing mass or maximizing stiffness. It also allows for the defining of the factor of safety. As this is an aerospace application, mass reduction is critical and therefore selected. A safety factor of 1.25 is inputted as defined in the problem statement.

As this is an aerospace application, mass reduction is critical and therefore selected. A safety factor of 1.25 is inputted as defined in the problem statement. Next is the manufacturing method, of which the following can be selected: unrestricted, additive, milling, 2-axis cutting, and die-casting. Additive manufacturing is chosen, from which build orientation, overhang angle, and minimum thickness can be selected. All six directions (X+, X-, Y+, Y-, Z+, Z-) were chosen to give the generative design study as much freedom as possible in its results. Overhang angle is left as its default of 45°. Its minimum thickness is entered as 1 mm. Lastly, the material that the results will be printed in is chosen to be titanium 6AI-4V. This will also cause the optical tracker to be made of titanium when it is stated to be made from aluminum 6061-T6 in the problem statement.

This will also cause the optical tracker to be made of titanium when it is stated to be made from aluminum 6061-T6 in the problem statement. While the Generative Design workspace allows for multiple materials to be included in the results, this is only valid for separate results; it cannot produce a result consisting of two different materials. Afterwards, the setup is complete, and the generative design study is ready to be converted into a CAD model for further testing.2087Fig. 4: A successful converged result in generative designFigure 4 shows a successfully converged result output from generative design. Theseresult outputs are visually inspected for any material discontinuities and material deposition abnormalities. If the results fail the visual inspection, they are discarded. If the results pass the visual inspection, they are converted into CAD models for further thermal testing.

If the results fail the visual inspection, they are discarded. If the results pass the visual inspection, they are converted into CAD models for further thermal testing. Successful models are explored below in the Results section.ResultsBased on the input parameters and the design space, Fusion 360 can produce multiple results using generative design. This is done in an iterative manner, with each iteration removing more material the software deems unnecessary. This is done until it produces a design that meets the defined parameters but with as little mass as possible. Once finished iterating, two types of results will be generated: converged and completed. Converged results means that the result fulfills all the inputted parameters, while completed results do not. Converged results are the only results that are of interest.Fig.

Converged results means that the result fulfills all the inputted parameters, while completed results do not. Converged results are the only results that are of interest.Fig. 5: a) Result from typical mounting pad geometry, b) Result from altered base mounting pad geometryConverged results are then converted into CAD files for further analysis. For instance, Figure 5a) shows an example result with “typical” pad geometry. Similarly, Figure 5b) shows a2088result with “atypical” pad geometry because of the creative liberties taken with the lower mounting pad geometry. The ability to experiment with lower mounting pad geometry allows for a tremendous range of design possibilities for further thermal testing.Thermal analysis1) Thermal loading scenario inputsThermoelastic deviation analysis begins with the application of thermal loads to a completed bracket model from generative design.

The thermal loads are referred to as Condition 1 and Condition 2; Condition 1 is defined as a 50℃ load applied to the tracker model and a 10℃ load applied to the base of the lower mounting pads, and Condition 2 is defined as a 10℃ load applied to the tracker model and a 50℃ load applied to the base of the lower mounting pads. The lower mounting pads are constrained in space during FEA, since the pads on a physical model would be fixed on a mounting surface. Because the model is expected to function in the vacuum of space, convective cooling is neglected, and the primary means of energy transfer is heat loss due to radiation. The background radiation temperature of space, -270℃, is used as the ambient temperature value. Note that the Autodesk thermal simulation environment uses degrees Celsius instead of Kelvin for radiation inputs. The selected material, Ti-6Al-4V, is applied to the structure, along with its thermal properties.

The selected material, Ti-6Al-4V, is applied to the structure, along with its thermal properties. See Figure 6 for a sample bracket with optical tracker model attached.Fig. 6: Finished bracket with tracker modelAs stated previously, the tracker model is necessary for subsequent calculations to determine the angular displacement of the centerline axis of the tracker instrument lens. The coordinates of the center of the tracker model are used to draw a vector line in space. This vector2089line is analyzed before and after thermoelastic deformation, and the total angular distance between the undeformed and deformed axis lines is determined.2) Test case resultsThe results from each thermal loading case are analyzed inside the Fusion 360 workspace. The model is analyzed for its displacement in the x, y, and z directions with respect to its original position in space. A typical case result is shown below in Figure 7.Fig.

The model is analyzed for its displacement in the x, y, and z directions with respect to its original position in space. A typical case result is shown below in Figure 7.Fig. 7: A typical displacement analysis result with plane of symmetry shownDue to the lateral symmetry of the bracket model, the bracket does not deform whatsoever in the lateral direction, which dramatically reduces the amount of total thermoelastic deviation under full thermal load. For these models, the lateral direction is defined as the x direction, and the plane of lateral symmetry is defined as the y-z plane. Lateral symmetry is advantageous for reducing thermoelastic deviation because the total amount of deviation in the x direction is mirrored about the plane of symmetry. The internal forces generated by the thermal loads are equal and opposite, resulting in a complete cancellation of any lateral deviation.

The internal forces generated by the thermal loads are equal and opposite, resulting in a complete cancellation of any lateral deviation. This results in thermoelastic deviation that is constrained to the y-z plane, and therefore a dramatic reduction in the total amount of thermoelastic deviation in comparison with an asymmetric model. Deviation results are collected for the x, y, and z directions, as well as total linear displacement. Deviation is calculated using the following formula:𝐴𝐴 ⋅ 𝐵𝐵 |𝐴𝐴| ∗ |𝐵𝐵|𝐶𝐶𝐶𝐶𝐶𝐶(𝜃𝜃) =where θ is the angle of deviation, A represents the original centerline through the tracker model, and B represents the centerline of the tracker after thermoelastic deformation. Vectors A and B are constructed by connecting coordinate points at the geometric center of the tracker model front and rear surfaces before and after deformation. The angle of deviation, θ, is then found and compared2090𝛳𝛳with the maximum amount of allowable deviation.

The angle of deviation, θ, is then found and compared2090𝛳𝛳with the maximum amount of allowable deviation. If the angle of deviation exceeds that of the design requirement, the design model is rejected as a possible candidate.3) Optimal ResultsA wide range of resulting structures have thus far been generated in Fusion 360. These results have dramatically different geometries and different thermoelastic behavior. While this research endeavor is still currently ongoing, early results are promising. Of the generated bracket design models, the design with the lowest thermoelastic deviation had a total maximum angular deviation of 0.0039°. See Figure 8 for bracket model. Although this deviation falls outside of the target deviation angle of 0.001°, it is a vast improvement over earlier models which deviated as much as 0.010° or higher from the original centerline position.

The performance of this bracket model, as well as the continuous refinement of each subsequent bracket model iteration, demonstrates the potential for this method to be used as a technique for the development of spacefaring structural components.Fig. 8: Best performing bracket model a) with tracker, and b) without trackerDiscussionThe implications of this research are compelling; it provides an alternative design and manufacturing methodology for the purposes of controlling thermoelastic deformation rather than eliminating it. It also capitalizes on the large number of intricate structural models that can be created quickly through automation, as well as the cost-reducing capabilities of using metal additive manufacturing for component fabrication. An additional benefit of the usage of metal additive manufacturing is the ease with which prototypes of a model can be made.

An additional benefit of the usage of metal additive manufacturing is the ease with which prototypes of a model can be made. Since the fabrication process with metal additive manufacturing is largely automated, bracket models can be fabricated and tested relatively quickly. The scope of this research involves testing each generated bracket model using the thermal FEA capabilities of Fusion 360. A proposed method to further verify the performance of generative design outcomes is the physical testing of prototypes manufactured using metal additive manufacturing. Testing methods for assessing the pointing accuracy of large earth-based telescope systems, such as the Large Synoptic Survey Telescope (LSST), involve the usage of laser trackers and a series of fixed calibration points [14].

Another2091optical system verification method proposed by a team at the Johns Hopkins University Applied Physics Laboratory involves the usage of a projected light dome to simulate a starry environment for the testing of optical systems [15]. Even further, an engineering team at the University of Bristol successfully trained an artificial neural network to predict the mechanical properties of generative design structures fabricated with filament deposition modeling (FDM), further strengthening the ties between neural network capabilities, generative design, and additive manufacturing [17].

Similar methods may be employed to test the accuracy and performance of prototype bracket models such as the ones generated in this study.It should be stated that the design and manufacturing methodology outlined in this paper is not meant to completely replace traditional engineering and manufacturing techniques for space applications; instead, this paper proposes the usage of generative design as one potential option for the design of spacefaring structural components. Previous research has gone into materials engineering for the development of thermally optimal materials to address the problem of thermoelastic deformation in space. This research endeavor merely aims to propose an alternative solution for the problem of thermoelastic deformation by controlling the deformation, instead of reducing it, through automated design. The appropriate design and manufacturing methodology should be chosen carefully after a cost-benefit analysis is performed.

The appropriate design and manufacturing methodology should be chosen carefully after a cost-benefit analysis is performed. For example, the usage of exotic materials such as HB-Cesic may be more appropriate for a spacefaring component, even with traditional manufacturing techniques. However, the usage of generative design with additive metal manufacturing opens the door for a wide variety of novel component designs that may be more suitable for a given application. Specifically, automation as a design tool allows for a tremendous number of highly intricate designs to be created in a very short time frame. A large number of unique designs can be generated fairly quickly, resulting in a reduction in lead times and cost of labor as a result; this reduction in labor cost alone makes automation an attractive option for the space industry.

According to the International Futures Program (IFP), the commercial space industry is projected to grow between 18-40% between 2004 and 2030, while entities such as Morgan Stanley estimate that the space industry will have a collective net worth of over $1 trillion by 2040 [13]. These figures suggest that the space industry may need to fully adopt automation as a design tool in order to meet future productivity needs. Generative and parametric design techniques have been shown to be potential solutions to bridge the growing productivity gap between the building industry and the manufacturing industry by facilitating design alternatives, improving information management, and fostering design efficiency [16]. Furthermore, it should be stated that automation will not replace engineers in the design process, but instead augment the abilities and boost the productivity of individual engineers.

Furthermore, it should be stated that automation will not replace engineers in the design process, but instead augment the abilities and boost the productivity of individual engineers. Engineers are still responsible for the definition of component usage cases and boundary conditions for design components and will still be required to evaluate the results from automation software.Acknowledgements This research was supported by the U.S. NASA MUREP Institutional Research Opportunity (MIRO) program (80NSSC19M0200), NASA MUREP High Volume (80NSSC22M0132), and partially carried out at the Jet Propulsion Laboratory, California Institute of Technology, under a contract with the National Aeronautics and Space Administration (80NM0018D0004).References20921. Goueffon, Y., Aldebert, G., Mabru, C., Arurault, L., Tonon, C. and Guigue, P., 2012. Flaking of black anodic films in space environment: Ageing and numerical simulation. Mechanics of Materials, 45, pp.72-82.2.

GENERIC SYSTEM FOR HOMOGENOUS SLS STEEL MATERIALSKlas Boivie Woxéncentrum, KTH, The Royal Institute of Technology SwedenAbstract A generic approach to a variety if different steel materials for SLS based application hasbeen examined. This approach is based on a base steel alloy powder mixed with a powder blend of much smaller particles. The powder blend is designed to both provide the highest possible density in the powder mass and melting material components for liquid phase sintering to full density. Furthermore, the liquid phase components in the powder blend are composed to both serve as a metallic binder for the green bodies and maintain the possibility to completely dissolve into the base material, forming a homogenous alloy. A powder blend with stainless steel base material was composed and tested and taken through the different process steps.

A powder blend with stainless steel base material was composed and tested and taken through the different process steps. While the formation of green bodies had limited success, to a large extent due to limitations in the process equipment, the sintering and diffusion behaviour showed promising results, both in respect to acquired densities, and homogeneity of the material.1. IntroductionWhile conventional polymer based SFF has established a position as an important strategictool for the manufacturing of prototypes, and is gaining momentum as a competitive direct manufacturing process for parts with complex or unique shapes, a similar development for the metallic applications still seems distant. The most important reasons for this may be found in the combination of the inherent constrains within the different approaches to metallic SFF processes, and the limitations in the selections of materials, available for SFF applications [1].

Therefore, developments in theses areas could significantly improve the capacity for new applications of SFF technology. Since steel materials are established for a very wide set of applications, with a long tradition as the engineering materials of choice, most potential users are likely to have steel materials as a major point of reference. Homogenous, SFF applicable steel materials, with similar properties as conventional steels, would therefore not only improve the capacity for SFF manufactured parts, they would also bring the technology closer to the users’ point of reference and thereby make it appear less alien to the potential user.However, for such a material development to have any significant impact, it should preferably by applicable to a SFF system with a wide distribution on the market.

This is possible with a procedure beginning with the formation of porous steel powder green bodies by the use of SLS, followed by metallic consolidation and homogenisation by liquid phase sintering in a subsequent furnace process. A generic approach to this procedure, where most variables can be kept constant, -except for those characteristic for the desired outcome, would not only offer SFF products of a homogenous steel material to a wider market, it would also facilitate the development of a much larger number of steel materials for a wide set of applications. (The term ”homogenous” is in this text referring to solid non-porous steel alloy, with phase structure corresponding to as-cast material of the specific alloy composition, with respect to heat632treatments.

This term is used to mark the difference from the composite or highly segregated materials used in some commercial application of metals SFF.)The whole procedure, from powder composition, application and processing in the SLSequipment, to the sintering and consolidation during the furnace treatment, is a vast area with a large number of variables, which could be used in many possible different combinations. Considering the endless possible variations, it is hardly possible to scientifically prove that such a material system and procedure is impossible. However, since this hypothesis is concerning technical matters, any successful attempt on any part of the procedure, would indicate that the whole procedure is possible, which is one principal object of this research, Another object is to find witch variables are critical to control, and how they affect the outcome of the process. In respect of this, the style of this work will be investigative and argumentative.2.

In respect of this, the style of this work will be investigative and argumentative.2. Background The formation of parts made of homogenous steel alloy, from SLS, with a subsequentfurnace procedure, is a very complex project were a number of variables may be of significant importance. This has been the subject for a series of investigations in the past [2], [3], [4], and [5]. This present work is based on conclusions and acquired knowledge from the preceding investigations as well as the application of established knowledge in conventional powder metallurgy. The approach can be separated into three distinct steps: powder composition, formation of green bodies, and the furnace process, were the limitations and results of each step, affects the conditions for the others.2.1. Powder compositionIf one step should be singled out as being the most critical, it should be the powder composition.

Powder compositionIf one step should be singled out as being the most critical, it should be the powder composition. Obviously, the powder composition determines the alloy composition of the finished material, but it is also fundamental for every part of each process step, from spreading and application of powder in thin layers, to the sintering into full density and homogenisation of the finished product.The principal idea of the present process approach, is to combine a base material powder, which’s alloy composition will dominate the finished material, with lower melting temperature components, that will serve as a liquid phase during sintering and ideally completely dissolve into the base material, forming the desired homogenous steel alloy.

As a most demanding example, a stainless steel, Anval 2205, has been selected as base material for this present investigation, however the principle as such, is generic and could be applied to a vast number of different steel materials.For liquid phase, previous work has exclusively used copper, [2], [4] and [5], partly forbeing established in conventional powder metallurgy, as a liquid phase with good sintering properties for iron based alloys, but also because it does improve some desirable properties, such as corrosion resistance, strength and toughness, [6].However, the amount of copper that can dissolve into the base material is limited, and forsuccessful liquid phase sintering into full density, with a limited amount of liquid phase, it is important to have the powder mass as dense as possible. This has been experienced and633addressed in the past [2], [4] and [5].

This has been experienced and633addressed in the past [2], [4] and [5]. It was found that using a suitable combination of size ranges and particle shapes in the correct proportions, will improve the density of the powder layers applied in the SLS equipment, [3]. Since the base material is expected to remain stable, and maintain the geometrical shape of the part during the furnace process, the base material should obviously be the largest fraction and also largest sized particles in the powder blend. Additional components, such as the melting material, should, if possible be kept at particle sizes that will fit into the voids between the larger particles.This approach has been tried with -44mm Anval 2205 powder, copper powder of thesmallest particle size available, and with an additional amount of small particle size carbonyl iron powder for further densification, [5].

However the high inner friction, and the unexpected unfavourable size range relation between base powder and copper in the powder mass gave the powder blend poor particle packing properties. This, in combination with the low fraction of melting material for the liquid phase sintering (8%), made it impossible to achieve full density in the material. Therefore has this present investigation changed the size range for the base material to 44mm - 75mm, and added a second melting material: ferrophosphorus. Fig.1. & Tab.1, Fig.2. Fig.3. & Tab.2, Fig.4. & Tab.3.Fig.1. Anval2205. A stainless steel material, with the characteristic smooth shape of gas atomised metal powder. This alloy has been selected as a challenging material with interesting properties. Size analysis: ASTM-E11 sieve, 99 %< -75m.m, 9 %<-45m.m.CSi MnPSCrNi0.0190.701.170.0170.00222.35.17Mo 3.02Nb 0.00Cu 0.16Co 0.03V 0.07W 0.02N 0.19O 165Table 1.

The chemical composition of the stainless steel base material, Anval 2205, O in ppm.Fig.2. Copper powder, 99.9% Cu. According to vendor1-5 mm spherical shape, which in reality obviously is not the case. This discrepancy is however most likely due to the method of measuring and calculating size. In many cases are such small sized particle’s sized estimated by a measured value for particle surface that is recalculated to volume and diameter assuming spherical shape. For a sponge like powder such as this, this method would indicate a much smaller particles size than really is the case. For this investigation this defect will not disturb the size range composition, however, it will limit the density improvement.634Fig3. Carbonyl Iron SM: almost pure iron, with the spherical shape typical of carbonyl metal powder. In this size range (2-8 mm), the interparticle friction is associated with van der Waal’s and electrostatic forces, and is thus increasingly independent of particle shape.

In this size range (2-8 mm), the interparticle friction is associated with van der Waal’s and electrostatic forces, and is thus increasingly independent of particle shape. Nevertheless, the spherical shape still is the best possible considering the achievable packing properties.C Min. 99.0 Max. 0.2 Max. 0.8FeON Max. 0.1Table 2. The chemical composition of the filler material, Carbonyl SMFig.4. Ferrophosphorus powder, coarse and uneven. Particles. This has normally a negative effect on powder packing properties. However at this size range are these properties dominated by the powder mass’ internal friction, associated with van der Waal’s and electrostatic forces. For this investigation, surface smoothness of ferrophosphorus particles will be of minor importance. Fe BaseTi Max 1.50 Table 3. The chemical composition of ferrophosphorus powder. Physical properties (Sympatec): 10% below 2.61 mm, 50% below 8.25 mm, 90% below 15.23 mCSiCr Max.

Fe BaseTi Max 1.50 Table 3. The chemical composition of ferrophosphorus powder. Physical properties (Sympatec): 10% below 2.61 mm, 50% below 8.25 mm, 90% below 15.23 mCSiCr Max. 0.50Mn Max 0.50P 15.60.78 Max. 0.50Phosphorus, P, is a well-known additive to powder compositions to improve sintering properties by liquid phase sintering, though other positive side effects, such as improved tensile strength, elongation and hardness, with small additions of P, has been reported from conventional powder metallurgy, [7]. The best results were acquired with additions of 0.75 % - 0.80% P, and the added amount of ferrophosphorus will therefore be adjusted to yield this content of P to the finished material.2.2.

The best results were acquired with additions of 0.75 % - 0.80% P, and the added amount of ferrophosphorus will therefore be adjusted to yield this content of P to the finished material.2.2. Formation of green bodies This process is expected to be capable to deliver hollow parts and geometries with deepcavities, and since the part in all cases will be subjected to a significant shrinkage during liquid phase sintering, a supportive powder during the furnace process is not desirable. Therefore, organic binders, requiring a supportive powder during the binder burn out and the early stages of sintering, should be avoided. This issue has been addressed in previous work, [4] and [5], and it has been found that despite the low absorbance of energy from a carbon dioxide laser, copper can be melted during SLS processing and thus work as a metallic binder. The green bodies formed are however quite brittle, with especially weak bonds between the layers.

The green bodies formed are however quite brittle, with especially weak bonds between the layers. On the other hand, in this present investigation, there is possibility that the second melting phase, ferrophosphorus, also could melt at process conditions and improve this situation.635Since copper has such low absorbance for energy transmitted by carbon dioxide lasers [8], itis more likely that it is secondary heating that melts the copper. The laser energy is absorbed by the iron-based particles, which in turn heats the adjoining copper particles until melting and bonding occurs. If so, the increased particle size will require more energy to raise the temperature of the base material particles to the melting point of copper (1084.5 °C) and even more for ferrophosphorus, (melting point 1168.0 °C), if the same heating mechanism is active as in copper.

Since the equipment available for this investigation is a rather old Sinterstation 2000, 50 W CO2 laser, with a measured maximum output of 41.5 W, this may be a crucial limiting factor for the formation of green bodies. However, the formation of perfect green bodies is not the primary object of this investigation, but the establishment of a generic principle for a series of homogenous steel materials that could be achieved from a procedure consisting of SLS and a subsequent furnace process. For this object, sintering and homogenisation behaviour is more vital, and it is sufficient to show that green bodies can be formed by SLS, while quality is of less importance at this stage.2.3. Furnace process conditionsThe feasibility, as well as process time and temperatures for the furnace procedure, in order to achieve a homogenous, liquid phase sintered steel material, are determined by some important boundary conditions.

First of all, the process parameters must allow the liquid components (here: copper and ferrophosphorus) in the powder composition to melt and sinter the material to full density. Second, the process time and temperature must be sufficient for the liquid material to completely dissolve into the base material, as well as allowing the alloy components from the base steel to homogenously be distributed within whole part, -both by means of diffusion. Third, since the risk of evaporation of alloy components from the material during the vacuum process is controlled by the combined effect of temperature and pressure, there is an upper limit on the acceptable temperatures, -or a lower limit for pressures, at process conditions. Meanwhile the pressure in the process chamber must be kept low enough to allow desoxidation of the material during the process.

Meanwhile the pressure in the process chamber must be kept low enough to allow desoxidation of the material during the process. Fourth, the processing time must be as short as possible, partly to avoid unnecessary grain growth, but also to keep the process times at a competitive level.However, this complex situation can be simplified by applying standard settings for vacuumpressure, and thus leave time and temperatures as the process controlling variables. Further simplifications can be made by application of conclusions from previous work. It has been found that at standard vacuum settings, temperatures above 1200 °C will cause evaporation of chromium from stainless steel powder, (1233.6 °C for 22.4% Cr) [5]. Therefore 1200 °C can be considered a maximum process temperature, which, in combination with vacuum, will allow the desoxidation reaction between carbon present in the powder blend, and oxide on the particles’ surfaces.

This leaves process time as the critical variable to control.Since the copper’s only means for dissolving into the steel material under these process conditions is by self-diffusion, and the longest predictable diffusion distance for any material under these circumstances is from the surface to the centre of the largest particles, the diffusion of copper through the base material will be the determining factor of the process times. This simplification may overlook the closure of large pores, but the occurrence of those, as well as their size and distribution, is not entirely predictable, and must be experienced experimentally.636The previously mentioned investigation, [5], also found that 4 h would be sufficient processtime for 8 % copper to completely dissolve into -44mm base material steel powder. However, in this present investigation has the size range for the base material been changed the to 44mm - 75mm, and ferrophosphorus has been added as a second melting material.

However, in this present investigation has the size range for the base material been changed the to 44mm - 75mm, and ferrophosphorus has been added as a second melting material. The increased particle diameter gives an increased maximum diffusion distance, which, reasonably, would increase the required process times beyond practical and competitive limits. However, that investigation and the following estimation of process time are based on the assumption of the atomic configuration of g-iron, which would normally be the phase present for most steels at process temperatures, especially since the added copper is g-stabilising. But, in this case, ferrophosphorus is added as a second melting material, and this could vitally change the conditions for the diffusion of copper and other alloying components.Phosphorus is known to have a powerful stabilising effect for a-phase and, compared to for example copper, a rapid diffusion behaviour in both a-, and g-phase.

Therefore, the hypothesis is, that the phosphorus will diffuse into the base material of g-phase configuration, ahead of copper and other significant alloying components, and thus cause a phase transformation to a-phase. Since the atomic configuration of a-phase steel allows much higher rates of diffusion for copper than g-phase, this would probably make it possible to achieve homogeneous distribution of alloying- and former liquid components in the finished steel material within reasonable process times. This, in combination with the additional amount of liquid phase and increased powder packing densities could make it possible to achieve a homogenous steel material with no, or very low porosity, by SLS and a subsequent furnace processing.2.4. CalculationsThe time to dissolve the copper into the base material is controlled by the fractional saturation of copper in the base material particles.

CalculationsThe time to dissolve the copper into the base material is controlled by the fractional saturation of copper in the base material particles. Darken & Gurry [9] gives us the relationship:For diffusion into spherical shape:Diffusion constant,DD= gFe Cu=Dt=.005043.02 L(1 -295kJmol8.1expRT2333log( 1)F, [Eq. 1]) s2cm, [Eq. 2], [10].Fractional saturation,F=C m Cs-C C00, [Eq. 3], and L is the diffusion distance, here the radius of thebase material particles. From previous work [5], we know that at 1200 °C, 8% copper can theoretically be completely dissolved in spherical -44mm steel powder, in 3.85 hours. For reasons previously mentioned, a particle size range of 44mm - 75mm has been used for this present investigation. With 75mm powder particles, the diffusion distance is 0.00375 cm. Added copper, sC = 9%, which gives F = 0.889. Applied to [Eq.

With 75mm powder particles, the diffusion distance is 0.00375 cm. Added copper, sC = 9%, which gives F = 0.889. Applied to [Eq. 1], mC = 8%, at 1200 °C saturation of copper is DtD Fe Cug112 = L17.0. At 1200 °C is=21.6⋅10t =ﬁ36942.h, which is too long. However,with phosphorous induced phase transformation, diffusion is expected to take place in a-phase where,) s(1 -2cmkJmol250Fe CuDD[Eq. 4], [10]. Under these conditions6.8== aexpRTD Fe Cua=17.1⋅108 -t =ﬁ20432.s (!), which is much more reasonable.637However, these ideal conditions cannot be fully expected under real process conditions. Forexample, if the material has sintered to full density, the particles will not be spherical during the whole dissolving process. Complete homogeneity concerning pore size and orientation, or particle distribution within the green body can neither be expected.

Complete homogeneity concerning pore size and orientation, or particle distribution within the green body can neither be expected. Furthermore, the liquid phase will be a solution melted copper and ferrophosphorus, which means that the concentration of the different elements will vary during the process, and during the diffusion of copper, the balance between a-, and g-phase in the outer perimeter of each particle is difficult to predict, and therefore also the real diffusion rate. Considerable longer process times will therefore be used for this investigation, however the calculations above demonstrates that complete dissolving and diffusion to homogeneity is not unreasonable to expect with this material composition and process conditions.3. ExperimentalThe experimental part of this investigation follows the principles outlined in the backgroundsection. The powder blend is composed to yield maximum applied powder density using the selected powders.

The powder blend is composed to yield maximum applied powder density using the selected powders. Green bodies are formed using SLS equipment, and processed to full density in a vacuum furnace. Samples examined regarding porosity and have been further analysed regarding alloy composition and distribution using SEM equipment at the Department of Materials and Engineering, KTH.3.1. Powder compositionThe different selected powder materials have to be combined in the correct proportions toyield the expected results. The amounts of the different liquid components is determined by their respective solubility in the base material, and possible desirable effects on the material properties of the final alloy. Copper has a good solubility in iron-based alloys up to around 8 %, at process conditions, which settles the share of copper. Phosphorus has been reported to have most positive effect on material properties at 0.8 % P.

Phosphorus has been reported to have most positive effect on material properties at 0.8 % P. The added ferrophosphorus contains 15.6 % P, which makes an addition of 5.12 % yield the desirable amount of P in the final material. This combination would give a liquid share of 13.12 %, which is slightly less than conventional, [11]. However if the experiment should indicate that the liquid share is too small for pore closure, it is possible to increase the addition of ferrophosphorus.Besides getting the right amount of liquid phase, it is also important have the highestdensity possible in the powder mass.

Since the amount of small sized particles for the liquid phase has been determined by other boundary conditions, the remaining means to control this issue is the combination of large particle base material, Anval 2205, and a fill material powder with much smaller particle size, in this case Carbonyl Iron SM.The suitable proportions was found experimentally using the procedure described in [3], with the amounts of copper and ferrophosphorus kept constant at 8 % respective 5.12 %. The final powder composition can be seen in Tab.4. Fig.5, showing the powder blend gives an impression of particle size proportions and packing behaviour. The density of the applied powder layers in the SLS equipment was measured by the procedure described in [3], and was found to range between 65.8 % and 67.4, which seems promising for the following process steps.638Anval 2205Carbonyl SMCuPFe375.88 % 11.00 % 8.00 % 5.12 % Table 4. The composition of the powder blend.Fig.5.

The composition of the powder blend.Fig.5. The powder composition after mixing.3.2. Formation of green bodiesThe determined powder composition was exposed to SLS processing for several attempts with successively increased laser power and decreased scanning speeds before any usable green bodies could be formed. This occurred at the very limit of machine capacity. Apparently, the assumption concerning the need for increased energy due to the increase in base material particle size was correct. Still, green bodies were formed, however of poor strength, with very weak bonds and high porosity between the layers, and with a strong tendency for curling of the layers.

Still, green bodies were formed, however of poor strength, with very weak bonds and high porosity between the layers, and with a strong tendency for curling of the layers. It is a likely assumption that only copper melted to serve as a binder, and that this situation could improve, -in terms of improving green strength by melting the of ferrophosphorus, with the use of a more powerful laser, such as the 100 W laser common in more modern versions of SLS equipment from DTM/3Dsystems, or the even stronger lasers used by EOS.3.3. Furnace processing Samples were sintered in a conventional vacuum furnace, in accordance with conditionsdescribed in the calculations section, and the principles described in [12]. The parts were placed on an alumina bed, in a graphite crucible, with no supportive powder and loaded into the furnace. The temperature was raised to 1200 °C at a rate of 200 °C/h. held at that temperature for 4 h, when the temperature was lowered at the same rate.4.

The temperature was raised to 1200 °C at a rate of 200 °C/h. held at that temperature for 4 h, when the temperature was lowered at the same rate.4. Results and analysisThe samples had sintered to solid objects and undergone some shrinkage, Fig.6, notable isthe fact that the samples with less successful layer adhesion have had less shrinkage in X –Y –direction, instead, the more separated layers have had a predominant shrinkage in Z –direction, whereas the samples with more successful adhesion between the layers have shrunk more equally in all three dimensions. Furthermore, some sagging of large overhangs can be noted. Looking deeper into the material we find that there are some remaining pores, predominantly oriented in between powder layers, Fig.7. However none of these imperfections are unexpected, and could probably be traced to the less than perfect conditions during the formation of green bodies.639Fig.6. Sintered samples.

However none of these imperfections are unexpected, and could probably be traced to the less than perfect conditions during the formation of green bodies.639Fig.6. Sintered samples. Shrinkage and some sagging of overhangs are evident. Especially notable is the shrinkage of the cylindrical samples in the bottom line. The same STL file and laser power was used for all of them, however with different scanning speeds.Fig.7. Pore structure, close to the boundary between two layers. The porosity increases with decreasing layer adhesion due to increased curling of the layers at lower laser scanning speedMore interesting, still, is the analysis of material alloy and distribution of alloy elements inthe material. It shows that complete homogeneity has not been reached, which is quite reasonable, complete homogeneity hardly exist even in conventional steels. Certain elements naturally concentrate into certain phases, and nothing else could be expected under these circumstances.

Certain elements naturally concentrate into certain phases, and nothing else could be expected under these circumstances. But the distribution analysis as shown in Fig.8 and Fig.9, displays encouraging tendencies.Fig.8. First distribution analysis of important elements covering an area of 58.9 x 44.6 mm. Fairly even distribution throughout the part, however complete homogeneity has not been reached. As could be expected have some elements, such as Mo, had the tendency to concentrate to certain phases as is normal in conventional steels. Also interesting is the area of Cu and Si concentration in the lower left side corner. Apparently has this been an area where larger amounts of liquid phase has been gathered than could be dissolved in the surrounding base material.640Fig.9. The second distribution analysis of important elements covering an area of 51.0 x 42.5 mm. The tendencies from previous analysis appear to be confirmed.

The second distribution analysis of important elements covering an area of 51.0 x 42.5 mm. The tendencies from previous analysis appear to be confirmed. There are some areas where Cu and Si have been enriched, while P and Mo have been concentrated to certain phases.From an overall view, all elements are fairly even distributed over both areas analysed.There are however some points were certain elements (for example Cu & Si) are especially concentrated, and other elements, for example iron, is sparse. This could probably be traced back to the liquid phase during sintering. Silicon is highly soluble in both liquid copper and iron (such as in ferrophosphorus), and since phosphorus has the higher diffusion rate, iron will leave the liquid phase ahead of copper, thus the dissolved silicon will concentrate to the liquid copper. As pores are closing during the sintering process, remaining liquid material, in this case copper, will concentrate to the remaining larger pores.

As pores are closing during the sintering process, remaining liquid material, in this case copper, will concentrate to the remaining larger pores. With the larger amounts of copper amassed in a small region, it is hardly surprising that all copper will not dissolve into the surrounding material within the limited process time.6415. Discussion and conclusionsThe primary aim for this investigation was to explore a possible generic principle to buildparts of homogenous steel, using SLS equipment in combination with a subsequent furnace procedure. Despite the obvious shortcomings of the parts manufactured, it is not unreasonable to say that this possibility has been confirmed.

Despite the obvious shortcomings of the parts manufactured, it is not unreasonable to say that this possibility has been confirmed. Solid pieces of steel were formed from green parts made in the SLS machine, and the analysis of the distribution of the principal alloy elements shows, with some exceptions, a reasonable even dispersion.However the material selected for this investigation was a stainless steel, and apart from themost general material properties, such as tensile strength, hardness, and others, the corrosion resistance of the new alloy still has to be determined. This could be problematic, however, since there still are some areas with significant porosity or with excess concentration of, for example, copper. How much this would influence the corrosion resistance by crevice- or galvanic corrosion is unclear, but any measurements on the available pieces would not be conclusive either.

How much this would influence the corrosion resistance by crevice- or galvanic corrosion is unclear, but any measurements on the available pieces would not be conclusive either. As mentioned earlier, the areas with excess copper concentration probably originates in large pores, and since the porosity is primarily oriented between the applied powder layers, these could thus in many cases be deduced to poor adhesion of powder layers. Small pores could be addressed by increasing the fraction of liquid phase in the powder blend, but the large pores, which could end up filled with remaining copper, are more difficult and the best would be if they could be avoided altogether. One likely way to at least decrease their frequency could be to improve the adhesion between the layers.The most obvious method to improve layer adhesion would be to increase the laser power.

One likely way to at least decrease their frequency could be to improve the adhesion between the layers.The most obvious method to improve layer adhesion would be to increase the laser power. A more powerful laser, such as those installed in more modern SLS machines on the market by DTM/3D Systems and EOS would penetrate deeper into the powder layer and thereby increase the melting of material in between the layers. Furthermore, the higher laser effect would enable the melting of the ferrophosphorus during SLS, thus increasing the metal bonding and thereby further improve the quality of the green bodies. On the other hand more extensive metallic bonding will also increase the temperature induced tensions in the green bodies, though these tensions will be relieved during the furnace process. If needed for complete pore closure, the fraction of liquid phase (here the ferrophosphorus) could be increased.

If needed for complete pore closure, the fraction of liquid phase (here the ferrophosphorus) could be increased. However, a less desirable effect, of increased fraction liquid phase would most likely be an increase of the small sagging observed on the sintered parts, and thus loss of part geometry. On the other hand this behaviour could probably be moderated by stronger metallic bonds in the green bodies that comes with an increase in laser power.The full process of making solid homogenous steel objects from an SLS approach could not be verified by this investigation. However, the acquired results concerning sintering and diffusion are encouraging, and while the observations of the SLS behaviour are challenging, they do not contradict this possibility.642References 1. Boivie, K, ”Systematic view of SFF Metal Tooling Approaches” Proceedings of the 3rd Annual International Conference on Rapid Product Development, 2002, Bloemfontein, South Africa, 2002.2.

GENETIC ALGORITHM WITH DESIGN OF EXPERIMENTS APPROACH TO PREDICT THE OPTIMAL PROCESS PARAMETERS FOR FDM G. Arumaikkannu*, N. Uma Maheshwaraa*, S. Gowri* * Department of Manufacturing Engineering, College of Engineering Guindy, Anna University, Chennai – 600025, India.AbstractThis paper describes a Genetic Algorithm (GA) with Design of Experiments (DoE) approach to predict the optimized surface roughness and porosity characteristics of the parts produced using ABS material on stratasys FDM 2000 machine. The Mathematical Model (MM) was developed by using Response Surface Methodology (RSM). It is to predict and investigate the influence of selected process parameters namely slice thickness, road width, liquefier temperature and air gap and their interactions on the surface roughness and porosity. The developed MM is the fitness function in GA in order to find out the optimal sets of process parameters and to predict the corresponding surface quality characteristics.

The developed MM is the fitness function in GA in order to find out the optimal sets of process parameters and to predict the corresponding surface quality characteristics. These results have been validated and the experimental results after GA are found to be in conformance with the predicted process parameters.IntroductionRapid prototyping refers to a group of commercially available processes, which are used to create solid 3D parts from CAD, also known as Layered Manufacturing Techniques (LMT). Rapid Manufacturing uses LMTs for the direct manufacture of solid 3D products to be used by the end user either as parts of assemblies or as stand – alone products. They also discussed the advantages of applying rapid prototyping for direct manufacture in terms of various costs, lead times, freedom in product design and production design.

They also discussed the advantages of applying rapid prototyping for direct manufacture in terms of various costs, lead times, freedom in product design and production design. Rapid manufacturing uses layer manufacturing techniques for the direct manufacture of solid 3D products to be used by the end user either as parts of assemblies or as stand alone products. Small parts with high geometric complexity to be made in relatively small volume are the most suitable candidates for rapid manufacturing (Neil Hopkinson and Phil Dickens 2001).Rapid prototyping trends and developmentsRapid Prototyping and Manufacturing (RPM) is still in its infancy. The physical models made by most of these systems cannot be used as working parts, mostly due mostly to material and economic constraints. The major problems in the current RPM systems include: part accuracy, limited material variety and economical performance. In Stratasys Inc.

The major problems in the current RPM systems include: part accuracy, limited material variety and economical performance. In Stratasys Inc. conducted experimental research using various control parameters and materials on FDM systems. Data is presented to help the user to choose the appropriate material for specific applications. The in- depth research on FDM processes results in system refinements in accuracy, speed and surface finish (Xue Yan and Gu 1996).Anitha et al (2001) analyzed the effect of process variables such as layer thickness, road width and speed of deposition on the surface roughness of the components produced by the FDM process using Taguchi technique.

They found that layer thickness is the most significant factor and other factors road width and speed of deposition also contribute to surface roughness.150Response Surface Methodology (RSM)Response Surface Methodology (RSM) is a collection of mathematical and statistical techniques that are useful for the modeling and analysis of problem in which a response of interest is influenced by several variables and the objective is to optimize this response. RSM allows for better understanding of relations between the inputs and the response. In the case of FDM process the inputs refer to slice thickness, road width, liquefier temperature and air gap and the responses are porosity and surface roughness. The relationship can be written in the form of a polynomial function describing a surface such as in equation (1).

The relationship can be written in the form of a polynomial function describing a surface such as in equation (1). This equation has been used in this work to develop a second order response surface mathematical model.kkkkY=b0+∑Xb ii+∑Xb iii+∑ ∑XXb i ijji1 =i1 =i1 =1 j = j i ≠(1)As RP is moving towards rapid manufacturing, there is an increasing demand on obtaining good surface quality parts (Pulak M. Pandey et al 2003). Richard H. Crawford (1993) discussed processing of geometric data, computer based analysis and design for SFF manufacturing. The discussion of geometric processing issues focuses on accuracy and completeness of the input models and the algorithms required to process such models.Fused deposition modelingThe FDM method forms three-dimensional objects from computer generated solid or surface models like in a typical RP process.

As the first step towards understanding the behavior of process parameters in RP-FDM 2000, to develop a mathematical model to determine the most influential parameters affecting porosity and surface roughness during fabrication of ABS polymer. Since the RP parts will be employed as functional products, modeling of their performance under various processing conditions becomes necessary. Experiments were conducted on Stratasys FDM2000 machine using specific combination of levels of significant process parameters. Required numbers of experiments have been conducted with desired process parameter settings. For the study of the surface roughness, the roughness average (Ra) was taken as a parameter and measured with Taylor Hobson surtronic 3+ surface roughness tester.

The weight of the components was measured by electronic weighing machine and the porosity of the components fabricated is calculated.Mathematical model for porosityMathematical model developed to predict the percentage of porosity had an adjusted R-squared value of 0.9694 and it is shown in equation (2)Porosity (%) = - 5.85298 + 353.71408 × A - 35.23580 × B - 0.051890 × C +306.30239 × D -580.30875 × A2 + 31.66310 × B2 - 77.24794 × A × B - 304.96094× B ×D(2)Where, A-Slice Thickness B- Road Width C-Liquefier Temperature D- Air Gap151From the analysis of variance the sum of squares and F-statistic suggest that road width is the most significant of the process parameters and also it is involved in interactions with slice thickness and air gap.Response surface graphs for porosity Effect of slice thickness and road width on porosityFigure 1 shows the effect of slice thickness and road width on porosity at constant liquefier temperature and air gap.

As seen from the surface plot, with increase in road width and decrease in slice thickness, porosity decreases. The porosity is found to be less for minimum slice thickness of 0.178 mm and higher road width of 0.8 mm.DESIGN-EXPERT PlotPorosity X = A: ST Y = B: RWActual Factors C: LT = 270°C D: AG = -0.005 mm17.688714.6154%11.5422, y t i s o r o P8.46895.39560.2540.2350.980B: RW, mm0.8110.6420.4740.2160.197A: ST, mm0.3050.178Figure 1 Effect of slice thickness and road width on porosityVoids formed between two layers at the junction of the crosshatched areas will cause air entrapment, thereby increasing porosity. A lower value of slice thickness is chosen because it reduces these voids between two layers, thereby decreasing the porosity (Arumaikkannnu et al, 2004).Effect of road width and air gap on porosityFigure 2 shows response surface graph for effect of Road Width and Air Gap on Porosity at constant slice thickness and air gap.

The minimum porosity occurs for this interaction at 0.8 mm road width and air gap of -0.01 mm. A larger road width decreases the air gap, covers up the sub perimeter voids, hence porosity decreases.152DESIGN-EXPERT Plot Porosity X = B: RW Y = D: AGActual Factors A: ST = 0.216 mm C: LT = 270oC16.573614.316212.0588%9.8014, y t i s o r o P7.54390.0000.9800.0030.811D: AG, mm0.0050.642B: RW, mm0.0070.4740.0100.305Figure 2 Effect of road width and air gap on porosityEffect of liquefier temperature on porosityThe Porosity decreases linearly with increase in Liquefier Temperature. The porosity is larger when the model temperature is minimum. Liquefier temperature is the temperature at which the model material is melted. As the liquefier temperature increases, the fluidity increases, resulting in bulging and latter flattening of the road as it is being laid down (Bharath et al 2000).

As the liquefier temperature increases, the fluidity increases, resulting in bulging and latter flattening of the road as it is being laid down (Bharath et al 2000). These make the material flow easy to occupy the unfilled spaces and hence decreasing the porosity. The minimum porosity occurs at a temperature of 290°C. Slower cooling of molten material and higher liquefier temperature promotes topside flattening because of the lower viscosity which enhances the flow process.The defect between adjacent roads occurs as a void or simply as a weak interface depends on the degree of bonding between adjacent roads.

The bonding between adjacent roads depends on the temperature during the formation of the roads and on the lateral flow of material.Mathematical model for surface roughnessSurface Roughness (Ra), µm = 1167.276 + 1347.281 × A + 141.465 × B - 9.933 × C - 4649.567 × D + 0.021 × C2 - 4.770 × A × C + 23512.090 × A ×D -0.493 × B×C- 989.508 × B× D(3)Where A – Slice Thickness B – Road Width C– Liquefier Temperature D – Air GapThe simplified model representing the surface roughness Ra in µm, as seen in equation 3, had an adjusted R-squared value of 0.9425. The mathematical equation developed has shown that the surface finish is a parabolic function of the road width, slice thickness, air gap and liquefier temperature. From the analysis of variance the road width, liquefier temperature and slice thickness were found to be the most significant process parameters.

From the analysis of variance the road width, liquefier temperature and slice thickness were found to be the most significant process parameters. The interactions between slice thickness and liquefier temperature and slice thickness and air gap influence more153on the surface finish. In addition the interactions between road width with liquefier temperature and road width with air gap are also found to be significant for the same. Hence all the four process parameters have the influence on the surface quality of RP products.Response surface graphs Effect of slice thickness and liquefier temperature on surface roughnessFigure 3 shows response surface graph for the effect of slice thickness and liquefier temperature on surface roughness at constant road width and air gap. The temperature settings are set manually, and therefore, all build layer thickness must be fabricated at the same temperature configuration.

The temperature settings are set manually, and therefore, all build layer thickness must be fabricated at the same temperature configuration. Minimum surface roughness occurs at a liquefier temperature of 275°C and a slice thickness of 0.254 mm. The liquefier temperature which is the temperature at which the model material melts should be high enough so that the fluidity is more and therefore the melt can flow easily and fill up the voids. Also at higher temperature the road bulges leading to flattening of the top surface of the road. This leads to better surface finish.

Also at higher temperature the road bulges leading to flattening of the top surface of the road. This leads to better surface finish. But a very high liquefier temperature (290°C) is not recommended because at higher temperatures the fluidity is excess which deteriorates the build especially on the top surface due and resulting in poor surface finish.DESIGN-EXPERT Plot Surface roughness X = A: ST Y = C: LT, oC Actual Factors B: RW = 0.642 mm D: AG = -0.005 mm26.81022.972m µ19.134, s s e n h g u o r15.29611.458e c a f r u S290.000280.0000.2540.235270.0000.216C: LT, oC260.0000.197250.0000.178A: ST, mmFigure 3 Effect of slice thickness and liquefier temperature on surface roughnessEffect of slice thickness and air gap on surface roughnessFigure 4 shows the response surface for effect of slice thickness and air gap on surface roughness.

Here, the surface finish is found to improve at a higher slice thickness of 0.254 mm and minimum air gap of -0.01 mm. Even at zero air gap there will be physically some gap between adjacent roads which leads to poor surface finish due to unfilled space formed. More negative air gap ensures that there is no gap between adjacent roads which covers up the unfilled space and leads to better surface finish.

More negative air gap ensures that there is no gap between adjacent roads which covers up the unfilled space and leads to better surface finish. Much better inter road bonding is obtained using a negative air gap setting which minimizes surface roughness and improves surface finish.154DESIGN-EXPERT PlotSurface roughness X = A: ST Y = D: AGActual Factors B: RW = 0.642 mm C: LT = 270 oC22.29818.958m µ15.617, s s e n h g u o r12.2768.936e c a f r u S0.0000.0030.2350.254D: AG, mm0.0050.0070.1970.2160.0100.178A: ST, mmFigure 4 Effect of slice thickness and air gap on surface roughnessEffect of road width and model temperature on surface roughnessFigure 5 shows the response surface graph for effect of road width and liquefier temperature on surface roughness. Better surface finish is observed when a high liquefier temperature of 275°C and a minimum road width of 0.305 mm were set.

Better surface finish is observed when a high liquefier temperature of 275°C and a minimum road width of 0.305 mm were set. At minimum road width, the adjacent road laid down will be very near to the previous one which leads to more uniform and more number of profiles resulting in minimum surface roughness.

At minimum road width, the adjacent road laid down will be very near to the previous one which leads to more uniform and more number of profiles resulting in minimum surface roughness. Also the higher melting temperatures make the road bulge because of decrease in viscosity and hence flow and flatten thereby resulting in better surface finish of the part produced (Berins, 1991).DESIGN-EXPERT PlotSurface roughness X = B: RW Y = C: LTActual Factors A: ST = 0.216 mm D:AG = -0.005 mm32.60026.95421.309m µ15.663, s s e n h g u o r10.017e c af r u S290.000280.000270.0000.6420.8110.980260.0000.474C: LT, oC250.0000.305B: RW, mmFigure 5 Effect of road width and liquefier temperature on surface roughness155Effect of road width and air gap on surface roughnessFigure 6 shows the influence of road width and air gap on surface finish.

Better surface finish could be achieved at a more negative air gap of -0.01 mm and minimum road width of 0.305 mm. A negative air gap makes two roads overlap which fills inter road gap and also the peaks formed by the road widths will be nearer to each other which results in uniform surface finish. This conclusion is in agreement with that shown in Figure 5. Better surface finish could be obtained with higher slice thickness, minimum road width and air gap, and at temperature of 275°C.DESIGN-EXPERT Plot Surface roughness X = B: RW Y = D: AG Actual Factors A: ST = 0.216 mm C: LT = 270 oC21.79318.70515.617m µ12.529, s s e n h g u o r9.441e c a f r u S0.0000.0030.0050.0070.4740.6420.8110.980D: AG, mm0.0100.305B: RW, mmFigure 6 Effect of road width and air gap on surface roughnessGenetic algorithmMany practical problems can be modeled as optimization problems. In order to solve these optimization problems, many methods have been developed.

In order to solve these optimization problems, many methods have been developed. Inspired by the principles of biological evolution, many simulated evolutionary algorithms have been developed. Hence optimization of FDM process parameters has been taken into considerations in this research work and an efficient genetic algorithm has been used for optimization. The input process parameters are optimized using Real Coded Genetic Algorithm (GA) to obtain the required characteristics. For this problem a real-parameter GA with an efficient crossover operator known as the Simulated Binary Crossover (SBX) is employed. The real-parameter GA with SBX has self-adaptive properties, which have been shown to exist in evolution strategies and evolutionary programming. The population is a collection of chromosomes during any generation and the number of chromosomes in a population remains a constant.

The population is a collection of chromosomes during any generation and the number of chromosomes in a population remains a constant. In this research work a chromosome represents the input process parameters namely slice thickness, road width, air gap and liquefier temperature. Each chromosome in the population consists of input process parameters with different combinations of respective values within their limits.Chromosome representationChromosomeST .221 mm .523 mm -.05 mm 262° CRWAGLTTable 1.0 Chromosome representation156The chromosome is a collection of four variables each representing the four input parameters. The variable takes up a random value between the ranges specified for each input parameter.

The variable takes up a random value between the ranges specified for each input parameter. Table 1.0 shows the representation of chromosome with values.Evolution processThe selection mechanism works like the “guidance system” of the “evolution process” inGA as shown in Figure 7.0.Figure 7 Evolution processDue to the stochastic nature of responses, this operator may result in moving the search in a wrong direction if fitness values are obtained from one or few simulation replications. The genetic operators such as reproduction, crossover and mutation make the genetic search to search a wide state space. Before the selection of operator takes place the worst chromosomes in the population are removed and the dead chromosomes are replaced with the chromosomes produced by the selected genetic operator so that the size of the population is constant throughout the cycle. By maintaining the population size constant throughout the search space extends to a wide area.

By maintaining the population size constant throughout the search space extends to a wide area. For example:Population size = 200, Number of generation = 250, The total no of chromosomes generated = 200 × 250 = 50000.157Genetic operatorsThe number of individuals dying during a generation is dependent on the closeness ofindividuals to the solution. The genetic operators used in this optimization are discussed below.SBX (Simulated Binary Crossover) CrossoverThe crossover operator is the main genetic operator. The crossover operator is used where there is a need for hybrid chromosome in the population. The hybrid chromosomes are needed at instances when most of the chromosomes in the population are closer to the solution. When the ratio of the number of died to the population size is lesser than 0.3 then SBX Crossover operator is chosen. The chromosomes needed for the crossover are chosen from the remaining 70 % of the population.

The chromosomes needed for the crossover are chosen from the remaining 70 % of the population. The SBX crossover operator is µ, which is a random value selected between 0 and 1 during each cycle.Child1 Child2= =Parent1 × µ + Parent2 × µ +Parent2 × (1-µ) Parent1 × (1-µ)The order allocation for Child1 and Child2 for a particular run is shown in Table 2.0.STRWAGLTParent 1.205 mm.516 mm.01 mm 252 ° CSTRWAGLTParent 2.183 mm.495 mm.05 mm 265 ° CRandom value µ=0.36STRWAGLTChild 1.191.503 mm.036 mm 260 ° CSTRWAGLTChild 2.197 mm.508 mm.024 mm 256 ° CTable 2.0 SBX CrossoverThe slice thickness for child 1 calculated as STchild 1 ∴ STchild 1 STchild 1= = =STparent1 × µ + + .205× 0.36 .191 mmSTparent2 × (1-µ) .183 × (1-0.36)158Mutation operatorMutation operator helps the genetic search to move in a random direction thereby eliminating the problem of being stuck in local minima.

A simple way to achieve mutation would be to alter one or more genes. In genetic algorithms, mutation serves the crucial role of either (a) replacing the genes lost from the population during the selection process so that they can be tried in a new context or (b) providing the genes that were not present in the initial population. The optimal mutation rate used for this problem is 0.05. Table 3.0 shows the mutation operation on a chromosome, wherein the slice thickness value of the chromosome is randomly assigned a value within the limits.ST.183 mmRW.485 mmAG -.006 mmLT285 ° CST.196 mmRW.485 mmAG -.006 mmLT285 ° CTable 3.0 Mutation operationA random chromosome is chosen and the input process parameter is also randomly chosen and its value is randomly assigned within its limits. Mutation operator is chosen when the number of died chromosomes to population size ratio exceeds 0.70. This occurs only when the population’s average fitness drops to 60 %.

Mutation operator is chosen when the number of died chromosomes to population size ratio exceeds 0.70. This occurs only when the population’s average fitness drops to 60 %. To improve the average fitness of the population new domains in the state space must be analyzed. Hence the mutation operator is selected for inserting random chromosomes in the population.Reproduction operatorThe reproduction operator is applied to emphasize good solutions and eliminate bad solutions in a population, while keeping the population size constant. This is achieved by identifying good (usually above-average- 95 %) solutions in a population, making multiple copies of good solutions, and eliminating bad solutions from the population so that multiple copies of good solutions can be placed in the population.

The commonly used reproduction operator is the proportionate reproduction operator where a chromosome is selected for the mating pool with a probability that is proportional to its fitness.Fitness functionThe MM developed using RSM is used as the fitness function for the GA.

The responses such as surface roughness and porosity were measured and calculated. These values were compared with that of GA results and they were found to be within their accuracy limits as shown in table 6.Response Porosity (%) Surface Roughness (µm)Predicted(GA) Experimental Accuracy (%) 4.231 25.324.012 24.2394.82 95.69Table 6.0 Validation results for mathematical modelConclusionThis work identified four important process parameters such as slice thickness, road width, liquefier temperature and air gap that had significant effect on porosity and surface roughness of the parts produced by FDM. that surface roughness increased with increase in road width and air gap and decreased with increase in slice thickness.It could be observed160The road width is found to be the most effective parameter affecting the surface quality and porosity of the part produced. Increase in road width and temperature reduces the porosity.

Increase in road width and temperature reduces the porosity. The regression statistics proved that the mathematical models developed have a high precision. Hence, those models could be used to predict the surface quality in terms of porosity and surface roughness within the range of various significant process parameters with 95 % confidence level. The developed GA could be used to predict the optimal process (input) parameters as well as the output responses accurately.References[1] Anitha R., Arunachalam S. and Radhakrishnan P. (2001), ‘Critical parameters influencing the quality of prototypes in fused deposition modelling’, Journal of Materials Processing Technology, Vol.118, pp.385-388.[2] Arumaikkannu G., Gowri S. and Muralikrishna S.

ometric Challenges in Designing Parts for Machining using Wire-fed DEDD. M. Vaughan*, L. Meyer*, C. Masuo*, A. Nycz*, M. W. Noakes*, J. Vaughan*, A. Walters*, B. Carter*, R. Wallace*Manufacturing Science Division, Oak Ridge National Laboratory, Oak Ridge, TN 37831AbstractWire-fed DED using MIG welding systems allows for high deposition rates above 30lbs/hr, enabling much larger parts to be printed than would be possible on other DED systems. However, a drawback to this high deposition rate is a relatively low bead resolution on the printed part. Post-processing using machining is usually required on any mating surfaces printed using wire-fed DED. Problems such as residual stress in the build plate and printed part, underbuilding, and path interpolations can all lead to insufficient material deposition and deviation from the desired shape. These areas where the printed part varies from the model can leave defects on post-processed surfaces.

These areas where the printed part varies from the model can leave defects on post-processed surfaces. This paper will cover common geometry issues that can arise from wire-fed DED and design changes that can be made to ensure that the printed design contains the required material to achieve the finished part.IntroductionAdditive manufacturing enables complex parts to be produced layer by layer with minimal amount of waste material. Metal additive manufacturing usually relies on the liquification of feedstock in the form of wire or powder to deposit each layer in the form of beads. The energy input for this process usually involves lasers, electron beams, or plasma arcs to melt the metal. Among the various types of metal additive manufacturing techniques available, wire-arc additive is among the highest deposition rates available.

Among the various types of metal additive manufacturing techniques available, wire-arc additive is among the highest deposition rates available. Wire-arc additive manufacturing (WAAM) can usually be divided into two different forms: Gas Metal Arc Welding (GMAW) and Gas Tungsten Arc Welding (GTAW). GTAW additive uses a tungsten rod as the electrode for forming an arc and then the feedstock wire is fed into the arc from the side. The tungsten electrode allows for fine control of the plasma arc. The GMAW process involves using spooled feedstock wire as the electrode to form an arc as it is continuously fed into the melt pool. This method of DED has benefits over other methods as it allows for a non-directional tool which simplifies tool path planning. The feedstock itself is also readily available due to its use in traditional welding processes.

The feedstock itself is also readily available due to its use in traditional welding processes. This paper willNotice: This manuscript has been authored by UT-Battelle, LLC, under contract DE-AC05-00OR22725 with the US Department of Energy (DOE). The US government retains and the publisher, by accepting the article for publication, acknowledges that the US government retains a nonexclusive, paid-up, irrevocable, worldwide license to publish or reproduce the published form of this manuscript, or allow others to do so, for US government purposes. DOE will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan). 1412focus on GMAW processes to produce large scale metal parts. WAAM processes can have instantaneous deposition rates between 6-10 lb/hr (1, 2). Other DED processes such as blown powder additive have lower deposition rates closer to 1-7 lb/hr (1, 3).

WAAM processes can have instantaneous deposition rates between 6-10 lb/hr (1, 2). Other DED processes such as blown powder additive have lower deposition rates closer to 1-7 lb/hr (1, 3). This makes WAAM processes very useful for producing large scale parts on the order of 1m or larger that would otherwise be extremely time consuming to print using other methods.The tradeoff with the high deposition rateprovided by wire-arc WAAM is the relatively poor resolution bead deposition and high heat input. Powder bed systems are often the highest resolution systems with spot sizes in the micrometer range (4). Meanwhile, WAAM systems are usually much higher with bead sizes on the scale of several millimeters (5). Due to the large bead size produced using wire-arc systems, post processing in the form of machining is usually required for any mating surfaces on a printed part. Figure 1 shows the cross section of two short walls printed using wire-arc process.

Figure 1 shows the cross section of two short walls printed using wire-arc process. The large bead size creates similarly large deviations in the final surface of a printed part on the scale of millimeters (5). The resolution can be improved by decreasing the heat input into the weld but requires a corresponding decrease in deposition rate (6). Parts printed using WAAM should be designed to take the large resolution into consideration to allow for effective post processing. Additionally, the heat input from WAAM processes is significantly higher than other metal additive techniques. Residual stress and changing layer heights based on geometry also require consideration when designing parts.Figure 1: Cross section of wire-arc printed walls.Challenges in Wire-Arc Design for Post ProcessingThe following section will cover characteristics of wire-arc additivemanufacturing that can require design modifications when attempting to post-process parts.

All of the examples shown below were printed at the Manufacturing Demonstration Facility (MDF) within Oak Ridge National Laboratory. All MDF wire- arc systems use Lincoln Electric R450 welders and Autodrive 4R220 wire feeders1413along with a ABB IRB2600 or IRB4600 and Abicor Binzel ROBO water cooled torches. In the majority of cases 0.045 wire is used on 1” or ¼” mild steel baseplates.As-Printed Surface FinishThe first and most visible potential issue when designing parts for post processis the surface roughness of the finished parts. When generating G-code for a given part, the bead size is usually measured as its maximum width. However, when depositing multiple layers, the bead takes on a rounded shape, resulting in a waviness in the surface where the surface does not extend as far as the maximum width. Figure 2 shows a heatmap for the deviation from a best fit plane for a wall printed in 410SS. There is approximately 2mm of variation in surface geometry.

Figure 2 shows a heatmap for the deviation from a best fit plane for a wall printed in 410SS. There is approximately 2mm of variation in surface geometry. If G-code is generated for a part using the average width as the bead width, the part will be printed smaller than the desired geometry in some places. This error can lead to the as-printed surface remaining exposed after machining. On the right of Figure 2 is a 3D scan of a part printed in 410SS highlighting an area where low spots in the as-printed surface finish led to an unmachined surface on the side of the part. Mating surfaces which are not able to fully machined may cause the part to be nonfunctional. Additional machining can be attempted to improve the surface, but only if tolerances allow for it.Figure 2: Left: Heatmap deviation from average plan on a 410SS printed wall.

Additional machining can be attempted to improve the surface, but only if tolerances allow for it.Figure 2: Left: Heatmap deviation from average plan on a 410SS printed wall. Right: Insufficient surface material for machining process.Build Plate Residual StressAnother common issue when using WAAM processes is deformation in thebuild plate due to the relatively high heat input used. There are two main components to deformation of the build plate. The first is the heat input on the build plate itself. Deposition on the top surface of the build plate creates asymmetric stresses within the build plate as the top surface is melted and then cools. The melted baseplate material1414contracts and causes the edges of the build plate to warp upward. The second component is the impact of the material being deposited on the build plate to create the part. At MDF, mild steel is typical used for build plates.

The second component is the impact of the material being deposited on the build plate to create the part. At MDF, mild steel is typical used for build plates. When printing on these build plate in materials other than a mild steel there will also be a difference in coefficient of thermal expansion. This difference in expansion as both materials are heated adds to the forces placed on the build plate and can increase the amount of warpage seen on the build plate after printing. Figure 3 shows a test case demonstrating the effect and scale of build plate warpage. In the figure, a 609.6mm x 101.6mm x 25.4mm build plate has had 3 layers of material deposited on it. The printed geometry was approximately 36mm x 550mm x 11mm. The deformation seen can cause issues when trying to produce accurate parts with features near the build plate. Build plate warpage is also an issue when attempting to machine the part afterwards.

The deformation seen can cause issues when trying to produce accurate parts with features near the build plate. Build plate warpage is also an issue when attempting to machine the part afterwards. The build plate is often left attached to allow for fixturing during the machining process and to provide well defined corners to set a coordinate system in a CNC mill. If the build plate is warped, the plate may not function for these purposes and the bottom of the build plate may need to be machined to create a flat surface prior to machining of the part itself.Figure 3: 3D scan of build plate showing exaggerated view of deflection in build plate and deposited material.Part Residual StressIn addition to the potential for residual stresses causing deformation in the buildplate, the same problem can occur in the printed part itself in some cases. When the printed part is asymmetric, there is the potential for imbalances in the stresses within the part as it cools.

When the printed part is asymmetric, there is the potential for imbalances in the stresses within the part as it cools. Deformation in the part itself is typically only seen in tall, thin1415features made of materials with high coefficients of thermal expansion. Figure 4 shows one example of this deformation in a curved wall section printed in 316L stainless steel. The curved shape of the wall allowed for asymmetry in residual stress, resulting in visible movement between each layer deposition and on the overall part.Figure 4: Curved 316L wall showing (a) directly after deposition, and (b) 20 seconds after deposition.After each layer deposition the wall expands and contracts varying amountsbased on the heating and cooling rates for that layer. The result is a printed wall that is warped from the desired final geometry. Figure 5 shows the results of a 3D scan of the wall as a heat map of geometric deviation in reference to the desired final geometry.

Figure 5 shows the results of a 3D scan of the wall as a heat map of geometric deviation in reference to the desired final geometry. This build was paused halfway through to allow the build to fully cool before continuing, resulting in a visible shift in layers due to residual stress.1416Figure 5: Curved 316L wall showing visible shift due to part warpage halfway through the build.Overhang FeaturesWhen producing more complex parts using wire-arc systems, the ability toaccurately print parts with overhang features becomes more important. The level of overhang that can be produced varies between systems but is often around 20°. Since the material being deposited is in the liquid form, overhangs rely on the material solidifying before it can flow off the previous layer since at least some of the material will have no surface directly below it. However, as the overhang angle increases, the material being deposited will increasing flow over the edge of the overhang.

However, as the overhang angle increases, the material being deposited will increasing flow over the edge of the overhang. Figure 6 shows an example of this where overhangs of increasing angle were printed with the same designed overall height, but can be seen to be produced with consistently shorter heights as angle increases. This change in bead height is due to more of the material rolling over the edge of the overhang causes overhung features to build shorter relative to other features in a part. When a part then needs to be machined after printing, this may result in the overhung feature being located lower than anticipated and there will be insufficient material to properly machine the surfaces.1417Figure 6: Height comparison of 2 bead walls with varying overhang angles.UnderbuildingThe same mechanisms that cause overhanging features to build lower thanexpected can also impact builds without overhangs.

When depositing material, each bead tends to overlap with adjacent bead due to the surface tension of the material and the bead spacing designed in the path plan. As a result of this, each bead takes on an asymmetric profile that slopes away from the bead that it is deposited adjacent to. The overall profile of the combined beads remains flat except for the outermost bead which does not have an adjacent bead on one side. This causes the outermost perimeter of the part to under build as consecutive layers continue to slope downward at the edges of a part. Figure 7 shows two examples of perimeter underbuilding.Figure 7: Underbuilding occurring on the perimeter of parts.Underbuilding along the edges of parts can result in missing material on features and leave unmachined surfaces after post processing.

This issue also has the potential to propagate as the outer bead underbuilds enough to not provide support for the beads further into the part.1418Incompatibility Between Bead Geometry and PartWhen designing parts for wire-arc systems, the bead size can cause issues whentrying to produce parts with small features on them. The bead size can also cause geometric inaccuracies on large parts if the toolpath for the part is created incorrectly. In order to produce a solid part, it is critical to ensure that each deposited bead is approximately spaced at even distances from neighboring beads. An example of these problems is shown in Figure 8.

An example of these problems is shown in Figure 8. To achieve this, it is often necessary to design parts with dimensions in multiples of the correct bead spacing.Figure 8: Two example toolpaths that are improperly spaced for bead geometry.Designing Parts for Post ProcessingOverbuildingIn order to ensure that the printed part contains the required material to fullymachine or post process afterward, the most effective solution is to modify model of the desired final part to add excess material to compensate for the potential geometric issues that have been previously discussed. Figure 9 below shows a 3D scan of a printed part overlayed over the desired final model.

Figure 9 below shows a 3D scan of a printed part overlayed over the desired final model. The part was printed taller and thicker than the final part to compensate for any underbuilding effects and to ensure that the surface could be fully machined without leaving any as-printed surfaces behind.1419PrintedFinalFigure 9: 3D scan of a printed part overlayed over final machined CAD model.To avoid any potential build plate distortion effecting features on the printedpart, additional material should be extruded down from the bottom surface of the part to raise the printed part around 0.25” to 0.5” off of the initial deposition surface. This also allows for easier sectioning of the part from the build plate.Match Features to Bead Width IntegersIn addition to overbuilding the part, taking into consideration the geometry of the bead will avoid leaving unexpected gaps in the part due to bead that are unable to fit into small geometries.

Figure 10 below shows two examples of designing parts for wire arc additive. Figure 10a. shows an airfoil profile that has been modified to allow sufficient room for the bead geometry to fully fill in the part. The long, tapered end from Figure 8 has been expanded to allow for a perimeter and inset to be printed and the rounded front of the airfoil has been squared off to avoid any gaps inside the material. Figure 10b. shows the path for a rectangular section that has been designed before path planning to match the width of four beads to ensure that there is no mismatch between the model and path plan.1420(a)(b)Final Part2 X Bead WidthDesigned geometryFigure 10: Toolpath plans for geometries designed for wire arc additive.ConclusionIn conclusion, the issues discussed in this paper can be summarized with three key characteristics of wire arc additive manufacturing. The first deformation of the part and build plate due to residual stress.

The first deformation of the part and build plate due to residual stress. Deformation of the build plate will occur almost always, while part deformation typical only occurs on specific materials. Second, due to the nature of wire arc additive, the deposited material is always temporarily in a liquid state and can be influenced by gravity. This can result in undesirable geometry as the material flows out. Finally, the rough surface finish that is produced due to the large bead size should be considered and sufficient overbuilding should be designed in to allow for the surface to be fully machined. Taking these characteristics into consideration will help towards designing parts that can be properly post processed and implemented into a system.References1421[1] Ahn, Dong-Gyu. 2021. “Directed Energy Deposition (DED) Process: State of the Art.”International Journal of Precision Engineering and Manufacturing-Green Technology 8 (2): 703–42.

OMETRIC ELEMENT TEST TARGETS FOR VISUAL INFERENCE OF APRINTER’S DIMENSION LIMITATIONSS. Chang1, 2,*, H. Li1, N. Ostrout1, M. Jhuria3. S.A. Mottal1, and F. Sigg11College of Imaging Arts and Sciences, 2College of Science, and 3Kate Gleason College ofEngineeringRochester Institute of Technology, 69 Lomb Memorial Drive Rochester, Rochester, NY, 14623, USA.Corresponding Author E-mail: shu.chang@rit.eduAbstractAs technologies advance in the field of additive manufacturing (AM), it increases the demand in using test targets to quantitatively appraise the performance of AM processes and parts. This study presents a unique concept to address the dimensional and geometric viability of three- dimensional (3D) printers with test targets that are unique and complementary to those currently available. We have named these distinct designed artifacts as Geometric Element Test Targets (GETTs).

We have named these distinct designed artifacts as Geometric Element Test Targets (GETTs). The concept for the targets is to rely on positioning and spatial frequency of geometric shapes to induce failures that are indicative of the system’s dimensional limitations. A distinguishing characteristic is that the dimensional failures can be inspected visually. Systematic evaluations of the limitations can be further conducted through contact or non-contact measurements. The initial GETTs include three suites of test targets: line, angular and circular suites. We will illustrate this concept with samples produced with fused deposition modeling printers. The potential applications of GETTs include standardization, reference targets, in-line system control, and more.1.

The potential applications of GETTs include standardization, reference targets, in-line system control, and more.1 . I n t r o d u c t i o nAs technologies proliferate in additive manufacturing (AM), the competence to quantify the ability of fabrication processes to produce consistent and defect-free parts with intended dimensions has become an important and central requirement in the AM technology advancement. The AM community has recognized that there is a high demand for measurement science in order to achieve predictable and repeatable operations and quality products [1].Numerous test artifacts have been developed and used to find the accuracy of printed parts or to benchmark the performance of AM machines and have been well summarized in recent publications by S. Moylan et. al. [2 and references therein], [3 and references therein] and L. Yang and M.A. Anam [4 and references therein].

Moylan et. al. [2 and references therein], [3 and references therein] and L. Yang and M.A. Anam [4 and references therein]. These standard test artifacts focus on the dimensional and geometric accuracy of intentionally designed and positioned parts [2]-[4]. The test artifacts include basic geometric shapes such as cylinders, cubes, etc. (will refer to as fundamentals) as wellas super-positioned geometric shapes such as combined or free forms (will refer to as composites). The artifacts are strategically positioned with the intent to expose changes and faults in dimensional and geometric attributes [2]-[4]. The shapes are either solitary units or the repetition of same/different sizes.The measurement tools can be gauges, calipers, micrometers, coordinate measuring machine (CMM), profilometer (contact or non-contact), optical microscopes, etc. The methodology used are measurements of coordination points and then least square fitted to reference CAD models [2]- [4].

The methodology used are measurements of coordination points and then least square fitted to reference CAD models [2]- [4]. The measurements focus on both dimensional accuracies and geometric accuracies, which Yang and M.A. Anam [4] have summarized into six attributes: straightness, parallelism, perpendicularity, roundness, concentricity, and true positions (for pin and for z-plane).These test artifacts have been used to measure part accuracies as well as benchmarking different AM technologies. There has been a call for complementary targets that “in order to reveal more information about the process quality, the features of the standard test part for process development should be designed in such a way that slight deviations from optimal process parameters would result in significant defects” [4].The intention of this research is to address this demand and to present Rochester Institute of Technology’s Geometric Element Test Targets (GETTs).

GETTs assess the dimensional and geometric viability of three-dimensional printers and processes with a unique test target concept transpired from two-dimensional (2D) graphic printing. These GETTs focus on the 3D printer’s dimensional viability (the dimension which can no longer be produced because of the system uncertainty) and the geometric viability (the geometric shape which fails to maintain the original or the designed form and orientation). GETTs also provides the possibility to determine the process capability visually in addition to measurements through metrological instruments. The following sections will present a few selected concepts of 2D graphic targets, the translation of the 2D targets into GETTs, and our initial validation of the concept on GETTs.2.

The following sections will present a few selected concepts of 2D graphic targets, the translation of the 2D targets into GETTs, and our initial validation of the concept on GETTs.2 . T e s t A r t i f a c t s P r a c t i c e d i n T w o - D i m e n s i o n a l P r i n t i n gIn 2D graphic printing, it is common to use test targets to stress a printing system’s attributes and elicit responses from different subsystems to assess the process’ performance [5]. Similar to AM test artifacts, the 2D test targets such as those produced by RIT are consist of fundamental (lines, dots, grey level, etc.) and composite (pictorial) elements.

Similar to AM test artifacts, the 2D test targets such as those produced by RIT are consist of fundamental (lines, dots, grey level, etc.) and composite (pictorial) elements. The fundamental targets are typically used for visual inspection or quantitative measurements of production/reproduction dimensional and geometrical viability while the composite images are for assessing human psycho-physiological responses through ranking [6].Many fundamental 2D graphic targets rely on inducing failures in the production of intended amplitudes of repetitive patterns at high spatial frequencies. The decrease in pattern sizes is to locate where the pattern will be distorted during the printing process. Some of the failures will have distinguished characteristics that can be identify visually (or through an eye loop) while others will require instrumental and statistical assessments.

Some of the failures will have distinguished characteristics that can be identify visually (or through an eye loop) while others will require instrumental and statistical assessments. 2D graphical targets are utilized to measure systematically these responses additionally in order to establish a transfer function between the digital representations and their corresponding print quality. An example of a system distortion is the laser beam size or charge spreading on the surface of a photoconductor in electrophotographic printers where the effect of addressability size is determined through the useof targets, and later used to predict output image quality.

The elemental targets typically consist text, curved and high spatial frequency lines, and halftones at different gray levels as exampled in [5], [6], [7], [8], [9].In this paper we will introduce three standard targets which are widely practiced in graphic and document reproductions, namely, a checkerboard, a ray, and a concentric circle pattern. These three patterns have characteristics that are deemed applicable to 3D printing and will be our initial focus for consideration on potential 3D targets.Figure 1 shows a checkerboard target which is a snap-shot from the corresponding region in the Synthetic Targets from RIT’s test targets publications [7]. The checkerboard pattern is used for assessing a printer’s dimensional resolution and dot gains. Figure 1 also contains a sequence of checkerboard patterns with increasing checker size, labelled as “1 dot by 1 dot” and so forth.

Figure 1 also contains a sequence of checkerboard patterns with increasing checker size, labelled as “1 dot by 1 dot” and so forth. These numerical figures represent the number of addressable dots in each checker square area (either on or off for black or white respectively). The checkerboard in Figure 1 was designed for a printer with an addressability of 600 dots-per-inch (dpi); that means “1 dot by 1 dot” checker square has an area of 42.3 µm x 42.3 µm. When used for dimensional examination, a printer’s resolution will be the “x dot by x dot” block where the pattern is clearly discernable visually or through eye loops. The checkerboard pattern can also be measured using a scanner or a microscope. The dot gains will be reflected in the over or under filling of the checker area through the instrumentation measurement.Designed for 600 dpi printer: Checker Size 42.3µ1 dot by 1 dot2 dots by 2 dots3 dots by 3 dots4 dots by 4 dotsFigure.1.

Example of RIT’s checker board resolution target that identifies the dot-by-dot dimension that a printer can produce and the amount of over or under filling of checker areas for process dot-gain or dot-loss.Depicted in Figure 2 are ray charts which are predominately utilized in resolving orientation- dependent dimensional feasibility. Figure 2 contains two simulated ray targets as contrast to illustrate the impact of the imaging spot aspect ratio (such as a laser spot) on the print quality [10]. The left ray target is simulated with a square spot and the right with elliptical spots oriented at 45° (with zero being the horizontal line). The spots are indicated at the lower left corner of each ray target respectively. The square spots on the left have resulted in a uniform distribution of lines shown in the left target. The directional printing spots, or the elliptical spots, give rise to an anisotropic distribution of ray densities.

The directional printing spots, or the elliptical spots, give rise to an anisotropic distribution of ray densities. The directional density distribution is revealed in the right ray target as emphasized by a darker banding oriented at 135° and perpendicular to the original spot orientation.Figure.2. Specifically designed features such as the ray lines draw out system responses, in this case the directional effect, from the printing processes.Figure 3 is a concentric circle target also known as the contrast-resolution test target [10]. The main features in Figure 3 are sets of concentric rings, with equally black and white spacing. Figure 3 presents the concentric rings at decreasing contrast (defined as the ratio of reflectance of two areas in an image) horizontally and increasing spacing frequency vertically.

Figure 3 presents the concentric rings at decreasing contrast (defined as the ratio of reflectance of two areas in an image) horizontally and increasing spacing frequency vertically. This circular target probes the capability of a printing system to reproduce fine detail of various orientations, frequencies, contrast and tonal values (similar to grey levels). Similar to the checkerboard resolution target, system effects can be reflected in the target, as demonstrated by the impact from difference halftone screens in the work of F. Sigg and D. Romano [8].Figure.3. Contrast-resolution test target [11] that identifies the cut-offs for both tonal contrast and spatial resolution for printers.The patterns illustrated in Figures 1-3 are typically of what printing professionals rely on to calibrate and optimize print systems as well as to identify process failures.

Quantifiable metrics for fundamental print patterns have been developed with standards to be used with microscopes and scanners. Standards are also established for composite or pictorial images to be qualitatively assessed with rankings. A point to notice, the fundamental patterns are what have been used for in-process quality assurance, process controls, and printer’s models [11].3 . T r a n s l a t i o n o f T e s t T a r g e t s f r o m 2 D t o 3 DThe objective of the translation of 2D test target into 3D aims at developing a simplified inspection tool and working standard to determine a printer’s capability. The intent is to complement the existing test artifacts which provide quantitative assessments of the produced geometric shapes and parts [2]-[4]. The research has a long-term goal of developing geometric element test targets to assess system responses of AM processes and to establish functional relationships between the responses and process parameters.

By exploiting and measuring systematic limitations, GETTs has the potential for in-line assessments and process controls.abdcFigure.4. Perspective view of checkerboard (a), flat ray (b), slanted ray (c), and concentric circle (d) GETTs respectively.We have constructed three suites of GETTs, line, angular, and circular suites in the likeness of their 2D counterparts. Our focus of this work is on the GETTs presented in Figure 4. Figure 4 shows a few examples of GETTs (presented counter-clockwise from top-left, a checkerboard (a), a flat ray (b), a slanted ray (c), and a concentric circle (d)). As illustrated in Figure 4, the contrast attribute in 2D graphics is replaced here by the height for both the checkerboard and concentric circle GETTs. The ray target, however, has been translated into flat and slanted rays.

The ray target, however, has been translated into flat and slanted rays. The slanted ray varies in z-dimension with a convergence to zero at the center of the ray.The design of GETTs follow the following guidelines:The fundamental units of GETTs are simple geometric elements, such as cubes, triangular wedges, cylinders, etc.The placements of elements are in reference to the process-related orientations. • GETTs contain repetitive elements with diminishing sizes and amplitudes (heights). • The dimensional range for GETTs design is based on the minimal addressability of the printing device under examination. When the accurate addressability is not known, standard measurement such as a metric or an English system is used considering the smallest element size to be less than the addressability but of the same order if possible.The dimensional changes might be continuous or discrete.

• The elements can be positioned in isolation, intercepted, or integrated with either the same or different elements.Elements can be raised or recessed with respect to a reference plane. • For integrated targets such as ray chart, create convergence.Expanding from 2D graphic target design has been attempted recently. B.H. Jerad et. al. have explored the conversion of the ray chart (also known as Siemens star) into a 3D artifact [12]. His adaptation is a star of a raised center and has diminishing star arms pointing outwards. The star was designed for quantitative measurement. Here the ray GETTs have raised edges with dimensional convergence to zero height at the center of the ray. This ray has the potential for both visual inspections and quantitative metrological instrumentations.In the same publication, B.H. Jerad et. al. [12] also create a “Manhattan” test artifact for examining dimensional accuracy throughout the 3D space.

Jerad et. al. [12] also create a “Manhattan” test artifact for examining dimensional accuracy throughout the 3D space. The “Manhattan” test artifact contains separated squares of the same areas at different heights for measurement at various positions in space, The GETTs checkerboard have the same height for each checkerboard unit. The checkerboard GETTs is to reveal a printer or a print process’ planar resolution.The existing artifacts also incorporate the technique in which a series of structures are decreasing in size and repeating geometric shapes [2]-[4], [13]. The repetitions are directed at covering measurement accuracy and precision for a range of dimensions [2]-[4], [13] and do not appear to be for graphic test target applications.4 . V a l i d a t i o n o f G E T T sWe have two Fused Deposition Modeling (FDM) engines, a MakerBot® Replicator 2X and a LulzBot TAZ 3 at our disposal.

V a l i d a t i o n o f G E T T sWe have two Fused Deposition Modeling (FDM) engines, a MakerBot® Replicator 2X and a LulzBot TAZ 3 at our disposal. The FDM process drives a thermoplastic filament to the heater and melts the thermoplastic to deliver the molten material through a nozzle [14]. The process thus deposits a thread of thermoplastic material to build the part line by line and then layer upon layer on a base platform. Although both of low resolutions and one technology, the two machines offer a comparison and contrast for the two systems.The GETTs designs were prepared using SolidWorks 3D modeling software. The design files are in STereoLithography (STL) format. We used Acrylonitrile Butadiene Styrene (ABS) in the validation experiment.For this study, both machines are set at default values with the assumption that these have been optimized by the original engine manufacturers (OEM). In this case both the printers have an extrusion temperature set at 230 °C.

In this case both the printers have an extrusion temperature set at 230 °C. The bed temperatures, however, are at different values of 110 °C and 85 °C for MakerBot and LulzBot respectively. We selected the infill density to be 10% for both the printers. As per OEMs’ specifications, the MakerBot has a nozzle opening of 0.4 mm and LulzBot of 0.35 mm.Photographs were taken with a Canon EOS 5D Mark Ⅲ camera. SOFV-1 light booth made by Graphic Technology provided the lighting for the photographs. Microscope images were captured using VHX-2000E series microscope made by Keyence© Inc at 20X magnification. The photographs are mostly used to demonstrate the feasibility of visual dimensional assessment. The images are captured by the microscope when better observation of specific characteristics are called for.

The images are captured by the microscope when better observation of specific characteristics are called for. The microscope has been set to minimal magnification to simulate the situation of inspection with eye-loops.The following Figures (5-9) concentrate on the concept of graphical patterns for distinguishing system attributes and features. Figure 5 depicts an example of the checkerboard pattern used to determine and compare the ability of two FDM printers to resolve square features in the horizontal plane. Exhibited in the left side of Figure 5 (Figure 5A) are photographs of four GETTTM checkerboard samples fabricated with the MakerBot. Positioned clockwise at decreasing checkerboard length and width, starting with the top-left photograph is the target that has checkers of 10 mm in length and width and 1 mm in height. The three checkerboards in Figure 5 have decreased dimensions of the checker’s length and width to 5 mm, 2.5 mm, and 1.25 mm respectively, all are at 1 mm height.

The three checkerboards in Figure 5 have decreased dimensions of the checker’s length and width to 5 mm, 2.5 mm, and 1.25 mm respectively, all are at 1 mm height. Exhibited in the right side of Figure 5 (Figure 5B) are the same GETTs produced with the LulzBot and presented in the same fashion.For MakerBot GETTs, the square shapes are well resolved in the 10 mm, 5 mm and 2.5 mm print samples, although there is some rounding at the corners of the 2.5 mm checkers. For the 1.25 mm checkerboard, however, the corners of the squares diminish and appear as circular protrusions instead of the angular squares. In comparison, the disappearance of corners come at 2.5 mm for the LulzBot GETTs. For the LulzBot samples, the checkerboard shape cannot be reproduced and lost its periodicity at 1.25 mm.According to manufacturer’s specifications, the MakerBot has a nozzle opening of 0.4 mm and the LulzBot has of 0.35 mm.

Figure 5 indicates that for both the machines, resolutions are far worse than the sizes of the nozzle openings or their addressability. Figure 5 shows that the MakerBot appears to have a superior square-resolving ability than that of LulzBot.A. MakerBotB. LulzBot10 mm5 mm10 mm5 mma)b)a)b)1.25 mm2.5 mm1.25 mm2.5 mmd)c)d)c)Figure.5. A sequence of decreasing sizes of checkerboard patterns shows resolution limitation difference of two FDM devices to produce squares. The device that produced the left group of GETTs, MakerBot, loses its ability at about 1.25 mm checker size. In contrast, the device that produced the right group of GETTs, LulzBot, fails at about 2.5 mm checker size.A set of flat ray samples are depicted in Figure 6 to illustrate their role in revealing the limit of printable line widths. All pie-shapes in the five ray GETTs are designed to have a length of 10 mm and a height of 1 mm.

All pie-shapes in the five ray GETTs are designed to have a length of 10 mm and a height of 1 mm. The pies have equal widths for both protruding features and separations in- between in each chart with the first on the left with the largest and the last the smallest partitions. The five ray GETTs in Figure 6 consequently have increasing number of pies: 6, 9, 12, 15, and 18 and are displayed from left to right respectively. The targets here are printed with the MakerBot printer.6 Pies9 Pies12 Pies15 Pies18 PiesFigure.6. A sequence of increasing number of pies in ray GETTs. The resulted center rings mark the areas that the printer has failed to print. These areas can be used to calculate the resolvable line width.The intent of the ray GETTs design is for the dimensions to converge to a singularity point at the center in this example. As this singularity is impossible to be produced by the printing system, a forbidden ring will emerge at the center of the ray.

As this singularity is impossible to be produced by the printing system, a forbidden ring will emerge at the center of the ray. The size of the forbidden ring reflects the line width producible by the FDM system. As indicated in Figure 6, the size of the forbidden ring grows in the proportion to the increase in the number of pies in the ray. This is because when the pie width gets narrower, the position of dimensional failure moves outward. If one may estimate the diameters of the rings presented in Figure 6 (say using the 15 pie ray chart here to get an about 4 mm diameter) and estimate the pie widths at which the failure occurs, a roughly 0.4 mm width can be obtained. With the OEM specified nozzle opening of 0.4 mm, the deposition process appears to transform a πx(0.4)2 mm2 circular area into a roughly rectangular area of 0.4 mm base and thereby an about 0.3 mm height.

Since the thermoplastic material will flow upon contact with the substrate, we expect to observe gains in widths of lines and loss in heights of lines.In complementary to the flat ray target, the slanted ray GETTs can be used to determine a FMD printer’s z-direction effective addressability. Figure 7 shows two six-pie ray GETTs whose designed dimensions are 2 mm in height at the end of the wedge and converging to zero height in the center with the pie length being 10 mm. The images here are capture with Keyence© VHX- 2000E series microscope by at a 20X magnification to better revealing the details at the center and the steps. As with the flat ray GETTs in Figure 6, both of them demonstrate the forbidden zones. In addition, the photographs show steps on the incline plane due to the nature of the FDM technology.A comparison of the two photographs reveals that for the same designed height MakerBot has generated 9 steps in the printed sample while LulzBot only 6 steps.

Measuring the resulting wedge heights indicate a reproduction of 2 mm with the MakerBot but only 1.6 mm with the LulzBot. Dividing the heights with their respective steps shows that MakerBot has a higher z-direction addressability of 0.22 mm/step than that of 0.26 mm/step from the LulzBot. In addition, only about half of the pies indicate distinguishable steps in the LulzBot target image. Furthermore, the LulzBot target image has a larger failed area at the center than the MakerBot target image, also depicting the inferior dimensional resolution of the LulzBot machine.MakerBot SampleLulzBot Sample9 Steps6 StepsFailed AreasFailed AreasFigure.7. Convergent slanted ray GETTs divulges FDM printers’ z-direction addressability through the number of steps on the slanted plane for a given height.Depicted in Figure 8 are two concentric circle GETTs produced using the MakerBot.

The two GETTs are designed to have 2 mm line protrusion heights but different spacing frequencies: 1 mm line width for GETT in Figure 8a and 0.5 mm line width for GETT in Figure 8b. A close inspection of the photographs reveals directional effects in Figure 8a, which shows distinct regions (as an example sections in Figure 8a separated by dashed lines) where the printed lines split while other regions the lines remain intact.As the line width is reduced (Figure 8b), the printed lines seem to be at the printer’s resolution and no longer split. However, there is noticeable unevenness in the line separation while other regions maintain the line spacing and width. As examples, two areas in Figure 8b are identified with arrows and marked as “Increased Gap” and “Decreased Gap”. In addition, a narrow section around 9 o’clock in Figure 8b exhibits a well-defined section which marks the start and end points of print process.a. 1 mm spacing frequencyb.

In addition, a narrow section around 9 o’clock in Figure 8b exhibits a well-defined section which marks the start and end points of print process.a. 1 mm spacing frequencyb. 0.5 mm spacing frequencyRegion of Line SplittingRegion of Line IntactFigure.8. Concentric circle targets to illustrate defects of FDM printing: line splitting, deviations in print paths, and visible joining of start and end points.We have assembled sequences of concentric circle targets similar to the 2D contrast-resolution target to present a “bird-eye” view of how variation in heights and widths of lines impacts the printer’s ability to reproduce the 3D circular element. In Figure 9, we replace the “contrasts” to line heights and present the same line width GETTs with diminishing line height horizontally from left to the right of the constructed assembly. We place the same height GETTs at increasing spatial frequency vertically from top to bottom of the assembly.

We place the same height GETTs at increasing spatial frequency vertically from top to bottom of the assembly. The concentric circle GETTs shown in Figure 9 are fabricated with the help of the MakerBot printer.By comparing the targets horizontally, we do not observe much difference in the GETTs production with changing line height. This is expected since the nature of FDM material deposition is to draw lines, in contrast when comparing GETTs vertically; the change in line spatial frequency has introduced a number of faults as already discussed for Figure 8. In addition, the vertical contrast reveals that when the spacing frequency is reduced to 0.25 mm, the concentric circle GETTs cannot be produced regardless of the line height.Start and End Region of Print ProcessIncreased GapDecreased Gap2 mm1 mm0.5 mm0.25 mmm m 1m m 5 . 0m m 5 2 . 0Figure.9. Height and frequency resolution concentric circle GETTs to detect dimensional failures and deviations.5.

0m m 5 2 . 0Figure.9. Height and frequency resolution concentric circle GETTs to detect dimensional failures and deviations.5 . P o t e n t i a l U s a g eIn 2D graphic and document production industry, targets are generally produced using high resolution printers and then used as references for both visual inspections and measurement standards for quality control. They are also widely used to calibrate printers on manufacturing consistencies within the same processes and to benchmark different processes. It is a common practice to have specific targets integrated into the printer’s process controls and feedbacks. The targets are also employed widely in the interactions with customers for meeting their requirements and for providing proofs for the quality of the products.

The targets are also employed widely in the interactions with customers for meeting their requirements and for providing proofs for the quality of the products. With research and exploration, GETTs has the potential to attain all these uses in the AM industry.As for research and technology development, GETTs offer a method to examine and optimize system performance as functions of process critical parameters and material specifications. Determinations of functional responses will enable tradeoffs between objectives and attributes in the engine architecture designs. Furthermore, with the establishment of a printer’s fundamental dimensional and geometric capability, there is a potential to predict and simulate build parts.6 . C o n c l u s i o nThis study presents a graphical concept to address the dimensional and geometric viability of 3D printers with test targets, named as GETTs that are complementary to those currently available.

This study has validated that dimensional failures can be observed and estimated visually with systematic target designs. The GETTs examples from three suites of test targets (line, angular and circular) presented in this study have demonstrated the feasibility in detecting geometric and dimensional errors. With further validation with high addressability machines and more machine- specific target designs, GETTs have the potential as a simplified print quality inspection tool, a working standard, a target for in-line system control, and for studying the system or subsystem responses to process parameters.The aims of this research are to develop geometric element test targets, measurement methodologies and to assess system responses of additive manufacturing (AM) processes.

By exploiting and measuring systematic limitations, the project seeks to determine transfer functions between digital representations and printed parts as well as to explore opportunities for use of the GETTs in in-line assessments and process controls.A c k n o w l e d g e m e n t sThis research is supported by the Melbert B. Cary Jr. endowment from the College of ImagingArts and Science at Rochester Institute of Technology.R e f e r e n c e s[1] NIST’s Measurement Science for Additive Manufacturing Program [online]. Availableat http://www.nist.gov/el/isd/sbm/msam.cfm.[2] Shawn Moylan1 , John Slotwinski , April Cooke , Kevin Jurrens , and M. AlkanDonmez, “An Additive Manufacturing Test Artifact,” in Journal of Research of the National Institute of Standards and Technology, Volume 119, pp. 429-459, 2014. [3] S. Moylan, J. Slotwinski, A. Cooke, K. Jurrens, and M. A.

ometric Modeling for Rapid Prototyping and Tool FabricationE. Levent GursozLee E. WeissFritz B. PrinzThe Engineering Design Research Center, and The Robotics Institute Carnegie Mellon UniversityAbstractThis paper describes the application of a non-manifold geometric mod eling environment, NOODLES, to a Rapid Tool Fabrication System. This system integrates stereolithography and thermal spray deposition into a CAD/CAM environment which includes design evaluation tools, robotic spraying, and computer-aided process planning. The level of integration and the number of different models in this system requires geometric representa tions that can be abstracted at several levels and that can be manipulated over several dimensions.

The models in our framework for design, analysis, and fabrication share a single comlnon unifying geometric representation implemented in NOODLES.1351IntroductionThe capability to ITlanufacture a wide variety of quality products in a timely and cost-effective response to market requirements is a key to global competitiveness. The opportunities for improving manufacturing technology range across the entire spectrum of industries, materials, and manufactur ing techniques. There is no single technological innovation which, by itself, will significantly improve productivity; rather it is a systems issue which involves rethinking many manufacturing activities. One such activity is the manufacture of tooling (Le., design, prototype, and fabrication) such as dies and molds required for the high-volume production methods that gener ate most of our manufactured products. Tooling manufacture is typically an expensive and time-consuming. process.

Tooling manufacture is typically an expensive and time-consuming. process. The reasons lie not only in the fabrication costs and time constraints imposed by conventional machining methods, but also in the organizational.framework. In most organizations, different groups employ different processes to design and manufacture tools and products. And the expertise in tool design and product design reside in different groups, impeding communications. between them. The representa tionql and physical models used in design,prototyping, and manufacturing, are often incompatible with one another, so transitions between the stages are time-cqnsuming and error-prone.• Products often make several complete cycles through design, prototyping, and fabrication before reaching produc tion. Thus, new product development or product modification implies a series of iterative changes for both product manufacturers and toolmakers.

Thus, new product development or product modification implies a series of iterative changes for both product manufacturers and toolmakers. For all these reasons, a rapid and smooth transition from product concept to production remains a challenge.This paper describes the geometric modeling aspects of a CAD/CAM tool manufacturing system to address this challenge for an injection mold In this system, both prototyping and tooling fabrication ing paradigm. are based upon compatible solid free-form fabrication processes, while the underlying geometric and process models share a common representational scheme.2 BackgroundSolid free form fabrication (SFF) processes [1] such as stereolithography, selective laser sintering, three-dimensional printing, and laminated object manufacturing quickly produce prototype parts. In these processes arbitrar ily complex shapes are built up incrementally in layers.

In these processes arbitrar ily complex shapes are built up incrementally in layers. The requirement to136decompose complex three-dimensional geometries into 2! dimensional layers make it difficult if not impossible to be driven by direct human operation. Furthermore, these processes usually require additional planning where hu man interaction would not be possible or appropriate. For instance, gen eration of scan vectors from the cross sections in stereolithography is one such operation. High-level control of free-form fabrication techniques thus require appropriate computer geometric representations for task description as well as for process planning.The Rapid Tool Manufacturing System [2] uses SFF prototyped parts in a sprayed metal tooling process to quickly build injection molds for manufac turing these parts in quantity. The system.-;urrently uses patterns produced with stereolithography apparatus (SLA). Thermal spraying is used to de posit metal onto the SLA patterns.

The system.-;urrently uses patterns produced with stereolithography apparatus (SLA). Thermal spraying is used to de posit metal onto the SLA patterns. By incrementally depositing multiple fused layers, a free-standing metal structure is formed by separating the metal shell from the plastic substrate. This shell can be used in the fabrica tion of a broad range of custom tooling including injection molds, forming dies, and EDM electrodes, by filling it's cavity with appropriate backing materials.To realize the maxhnum benefits of SLA and sprayed tooling, these pro cesses must be effectively integrated into the overall manufacturing system. Conventional net-shape manufacturing systems (i.f~. those which produce the individual parts used to assemble the finished products) are composed of separate organizations, each with their .own expertise, equipment, and CAD environments.

those which produce the individual parts used to assemble the finished products) are composed of separate organizations, each with their .own expertise, equipment, and CAD environments. The transfer of information between organizations is error-prone and inefficient since computer models must be transformed be tween each environment by additional translators. Independent CAD sys tems make the software difficult to manage. Further, the lack. of down stream process knowledge (i.e. understanding how parts are manufactured) in the design stage leads to time-consuming and costly iterations. Our ap proach seeks to integrate design, prototyping, and tool fabrication into a common organizational framework.

Our ap proach seeks to integrate design, prototyping, and tool fabrication into a common organizational framework. The keys to integration are to incor porate "design-for-manufacturability" into the design stage, to automate robotic spraying, and to use a unifying geometric data structure for part representations and process planning.Some of the steps in the tool manufacturing process and the computermodeling requirements for each are summarized below:Design - The user designing a part should be allowed to select the appropriate Inodeling description paradigm depending upon the im Inediate need.

For example, designs, at times, can best be synthesized137using constructive solidgeom.etry, or building solids up fromsets of surfaces, while at others, sweeping lower dimensional elements, such as curves and surfaces, into solid representations produce more satis factory results.Pattern generation - Spray patterns are generated from design mod els by automatically deriving parting lines and establishing parting surfaces. Surfaces are added to solid model representations to create shapes which represent the complements of the mold halves.SLA - Building patterns with SFF require valid solid representations. Thus, if the design originates in the surfaced forms, it is vital to have a capability which transforms enclosing surfaces into solid represen tations.

Thus, if the design originates in the surfaced forms, it is vital to have a capability which transforms enclosing surfaces into solid represen tations. The SLA process planner must convert solid models into an ordered set of 2~ dimensional cross sections (i.e., cross sections with an associated depth or thickness) and span these cross sections with ap propriate drawing vectors. This operation inherently involves working simultaneously in several dimensions since one generates planes from solid models, and then vectors, or line segments, from the planes.Thermal spraying Robotic spraying, driven by an off-line trajec tory planner, automates thermal spraying and facilitates process con trol by its consistent and tireless performance. Off-line trajectory plan ning based on design models does not require tedious teach-by-showing operations, and the incorporation of expert rules to formulate spray strategies produces better quality shells.

Trajectories for spraying can be generated by projecting grids onto the object models. Interfer ence of the spray stream with the object is assessed by intersecting lines representing the spray cone with the solid corresponding to the pattern. 3 NOODLES ModelingGeometric modeling can be performed at various levels, such as wire frame, surface, or solid modeling. Topological information is required for feature extraction. The previous examples suggest that all levels of modeling and complete topological relationships are required in the system. Although solid modeling approaches have the richest information, the representation of lower level elements such as lines and surfaces is not explicit. Furthermore,138operations provided within solid modeling approaches do not apply when non-solid elements are used. The ideal geometric modeling system should uniformly represent and operate on non-homogeneous (Le., mixed dimen sions) elements such as vertices, lines, surfaces, and solids.

The ideal geometric modeling system should uniformly represent and operate on non-homogeneous (Le., mixed dimen sions) elements such as vertices, lines, surfaces, and solids. NOODLES offers an environment where non-homogeneous elements are uniformly represented and permits Boolean operations between elements of any dimensionality [3]. In addition, the NOODLES representation is rich in topological information which relates adjacencies between all geometric modeling elements.NOODLES employs a non-manifold boundaryrepresentation.scheme to model geometric objects. Along with the geometry of the elements that make up the objects, the topology is also captured explicitly. Conventional bound ary representation schemes model solid objects by their enclosing shells. One significant constraint in such schemes is that the topology of the enclosing shells are assumed to be two-manifold. This assumption implies homeo morphism to an open disc in the neighborhood of every point on the shell.

This assumption implies homeo morphism to an open disc in the neighborhood of every point on the shell. Data structures based on the two-manifold assunlption cannot gracefully accommodate elements of lower dimensionality which do not contribute to the construction of shells. Since NOODLES is based on non-manifold (or non two-manifold) topology, objects such as points, wire-frames and stand alone surfaces which are in violation of disc homeomorphism can be grace fully represented along with solids. Furthermore, these non-homogeneous objects can interact with each other in set-theoretical operations.The capability to unite, intersect, or subtract models and elements of models is an essential compon.ent of any geometric modeling system, and is typically implemented as Boolean set operations. In NOODLES, the notion of Boolean operation is applied literally to point set topology.

In NOODLES, the notion of Boolean operation is applied literally to point set topology. The legal operands of any binary operators (union, intersection, and difference) can be of the form of any point set element, namely, Vertex, Edge, Face, Solid. As a matter of convenience, a Model represents the union of all fundamental elements, and a Pset represents any arbitrary collection of those elements.4 ExamplesSeveral examples are presented which demonstrate the power of such non-manifold modeling capability. First, if a design is represented by a set of surfaces, then it must be transformed into a solid representation in order to build it with SLA. When surfaces are introduced, NOODLES identifies and keeps track of enclosed volumes which thus define solids (Figure 1). An other example which uses non-homogeneous representations is the planning139of the layering process. The first step is to obtain the cross sections of the object.

An other example which uses non-homogeneous representations is the planning139of the layering process. The first step is to obtain the cross sections of the object. These sections are obtained from the Boolean intersection between the object and a stack of planar faces that are appropriately spaced. The result of this non-regular operation (Le., operations between entities of dif ferent dimensions) is a collection of cross sections (Figure 2). The vectors to be scanned by the laser are obtained by intersecting appropriate grids with the portions of the cross section. For example, the interior area of a cross section is is intersected with a cross hatch grid (Figure 3). The ob ject boundaries for the laser are quickly found frora the perimeters of the cross sections. Similarly, the grids for robotic path planning are defined by the perimeters of the intersection of the surface boundary of the object with mutually perpendicular sets of stacks of planar faces (Figure 4).

The robot trajectory is defined by these contours and by specifying the stand-off distance of the spray gun. The spray gun orientation is optimally oriented normal to the surface. Interference is assessed by intersecting lines which represent the planned trajectories of the particle spray with solid models of the pattern. When interference is detected, the spray gun is incrementally reoriented until a direct line-of-sight is achieved (Figure 5).5 ConclusionSince most SFF processes build solid objects by creating successive layers on top of each other, the most fundamental geometric operation is to obtain cross-sections from the solid models of the objects. It is very desirable to em ploy a uniform geometric modeling environment that can represent the solid objects and the cross-sections under the same representational formalism. This is important from the viewpoint of system integration and informa tion flow.

This is important from the viewpoint of system integration and informa tion flow. Under such an environment, the design process which results in a creation of a solid model of the artifact and the SFF process planning operations can be carried out using the same geometric modeling system. This eliminates the need to use several geometric modeling schemes and the task to translate the representation between different schemes. Futhermore, the representational instances of the solid object and the representational instances of the cross-sections can be informationally linked. Such linking would smoothly propagate the attributes of the solid object to the level of cross-sections. For example, The surface finish requirements specified for the selected surfaces on the solid object can easily be made available as an inherited attribute at the contours of the corressponding cross-sections.

This140information can then be used to adapt the process to maintain the specified surface finish.In addition to the uniformity in the geometric representation between the design and the plannining phases, uniformity in geometric modeling op erations across these phases is also desirable. The same geometric modeling operations that are used in the design phase can be applied in the process planning phase. The creation of the solid model in the design phase typi cally involve modeling operations such as booloans, sweeps and extrusions. In the planning phase, cross-sections can be conceived as boolean intersec tions between appropiately elevated planes and the solid objects.In this paper, benefits of a non-homogenous geometric modeling envi ronment in the context of a rapid tool fabrication system were discussed. The system integrates streolithograpy and thermal spraying processes to manufacture injection molding tools.

The system integrates streolithograpy and thermal spraying processes to manufacture injection molding tools. The geometric modeling environment, NOODLES, is employed at various stages of the fabrication process to rep resent and operate on the different geometric abstractions appropriate for each stage.6 Bibliography[1] Machlis, S. "How To Bring CAD To Life," Design News, March, 1990.[2] Weiss, L.E., Gursoz, E.L., Prinz, F.B., Fussell, Mahalingham, S., and Patrick, E.P., "A Rapid Tool Manufacturing System Based On Stereolithog raphy and Thermal Spraying", AS ME's Manufacturing Review, March, 1990.[3] Gursoz, E. Levent, Choi, Y., and Prinz, F.B., "Vertex-based Repre sentation of Non-manifold Boundaries," in Geometric Modeling For Product Engineering, M.J. Wozny, J.U. Turner and K. Preiss Editors, Elsevier Sci ence Publishers B. V.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperGEOMETRICAL ACCURACY OF HOLES AND CYLINDERS MANUFACTURED WITH FUSED DEPOSITION MODELINGF. Knoop, V. SchoeppnerDirect Manufacturing Research Center (DMRC), Paderborn University, Paderborn, Germany and Kunststofftechnik Paderborn (KTP), Paderborn University, Paderborn, GermanyAbstractA widely used Additive Manufacturing (AM) technology is Fused Deposition Modeling (FDM) to create prototypes and end-use parts with close-to-production thermoplastics. For their use as a final product, it is necessary that additively manufactured parts strictly adhere to the geometrical requirements of the technical drawing. In this paper, the holes and cylinders of the cylindrical elements are investigated in terms of achievable geometrical accuracy.

In this paper, the holes and cylinders of the cylindrical elements are investigated in terms of achievable geometrical accuracy. For this purpose, different test specimens that allow a measurement of inner and outer diameters from 3 to 80 mm were designed. All specimens were measured with a coordinate measuring machine (CMM) to evaluate deviations from the nominal dimension and form deviations. The measuring method includes a scanning of the surface to record the course of dimensional deviations over the diameter. Thus, it was possible to visualize how deviations on cylindrical elements manufactured in FDM occur. In order to counteract these deviations and to improve the dimensional accuracy, different shrink factors and filling patterns were investigated. Consequently, an improvement of the dimensional accuracy was achieved.IntroductionAdditive Manufacturing (AM) processes are characterized by a layer-by-layer material deposition so that no forming tool is necessary.

It is possible to produce very complex parts which are impossible or very expensive to manufacture by conventional manufacturing processes. The starting point for every AM process is a 3D CAD model from which the layer information is generated in the data preparation [2]. Fused Deposition Modeling (FDM) is a very widespread AM process and was introduced onto the market in 1991 by Stratasys. FDM is the first additive manufacturing process that works according to the basic principle of material extrusion. After expiry of the basic patent in 2009, many open source devices entered the market, which operate in a very similar manner. The material extrusion systems still represent the largest share of all AM machines on the market [1,2]. The fundamental principle of the FDM process is based on the extrusion of polymer material through a nozzle.

The fundamental principle of the FDM process is based on the extrusion of polymer material through a nozzle. A pinch roller feed system above the nozzle ensures that the filament material is pushed through the nozzle and that the necessary extrusion pressure is generated [4]. The necessary energy for melting the polymer is provided by electrical heating elements. Extruded material exits the nozzle and is deposited as a thin filament by the FDM head. The FDM head is moveable in the X- and Y-directions (see Figure 1). The thermoplastic material solidifies by heat conduction and thus by cooling, resulting in a solid bonding [3,4]. After completing a part layer, the build platform is lowered by the thickness of a layer. A support material is used to join parts to the build sheet and to support overhangs or cavities.

After completing a part layer, the build platform is lowered by the thickness of a layer. A support material is used to join parts to the build sheet and to support overhangs or cavities. The support material is applied by means of a second nozzle and can be removed either mechanically or chemically after the process [7].2757Raw MaterialFDM HeadModel MaterialSupport MaterialBuild PlatformZBuild SheetYXFigure 1: Schematic representation of the FDM processThe FDM process is characterized by the fact that close-to-series thermoplastics can be used [5]. In contrast to established plastic processing methods such as injection molding, the mechanical [6,7] and geometrical properties [8-10] as well as the surface roughness [11] of FDM parts are greatly dependent on the alignment within the process. Fused Deposition Modeling is used for the manufacturing of prototypes and end-use parts in individual and small series.

Fused Deposition Modeling is used for the manufacturing of prototypes and end-use parts in individual and small series. FDM parts, which are often deployed in the aircraft industry (e.g. interior or air ducts), require high and consistent part quality. Apart from the mechanical and surface properties, the part quality also includes geometrical accuracy [7]. For the widespread industrial manufacture of FDM end-use parts, the dimensional accuracy and the occurring deviations must be known, and the influencing factors must have been understood [8]. In technical drawings, a geometrically ideal or nominal shape is assumed. However, deviations always occur during the manufacturing process. These deviations can be classified according to different properties. Thus the deviations are differentiated by dimension, form, location and surface [12].State of the ArtIn many previously published studies, dimensional deviations of outer dimensions, manufactured with FDM, have been investigated.

In almost all cases this is a measure of length, which describes the distance between two parallel opposite surfaces. Therefore, the test specimens are often non-complex elements like cuboids. LIENEKE et al. [8], KNOOP et al. [9] and DAO et al. [13] investigated the influence of the nominal length on the occurring dimensional deviations. The result was an increase in dimensional deviations with increasing nominal length. The authors PARMAR and PARMAR [14] and SOOD et al. [10] examined the influence of different FDM process parameters on the dimensional accuracy of a geometrically constant specimen. They conclude in their publications that the accuracy along the length and width of the part is very much affected by shrinkage. The thickness of the part is mainly influenced by the layer thickness. In publications by MINETOLA et al. [15], JOHNSON et al. [16] or MAHESH et al.

The thickness of the part is mainly influenced by the layer thickness. In publications by MINETOLA et al. [15], JOHNSON et al. [16] or MAHESH et al. [17] benchmark parts are used, which combine many geometric elements such as spheres, cylinders or holes in one part. The publications by SUDIN et al. [18], TAGORE et al. [19] and BAKAR et al. [20] deal exclusively with cylindrical elements manufactured with FDM. In addition to external and internal dimensions, roundness has also been investigated as a form deviation. SUDIN et al. varied the diameter of the hole and cylinder between 5 and 20 mm in four steps. The specimens placed in the XY-plane were made on a Fortus 400mc with ABS and a slice height of 0.254 mm. The result of the investigations is that cylindrical elements cannot be manufactured as accurately as slots and that the deviations are often outside the manufacturer’s accuracy of ± 0.127 mm. In the2758publication of TAGORE et al.

In the2758publication of TAGORE et al. the roundness of cylinders and bores with two different diameters, which are orientated between 0° and 90° to the build plane, is considered. The measurement is carried out by means of a coordinate measuring machine (CMM) at different heights of the specimen but without an exact specification of the measuring method. BAKAR et al. manufactured five cylinders with different diameters between 10 and 15 mm without changing the alignment of the parts. The influence of the layer thickness and raster width on dimension and form deviations was determined by measuring the specimens by means of a CMM. In the above-mentioned publications, only a very small variation of the nominal diameters (up to a maximum of 20 mm) was performed so that the influence of different dimensions could not be determined. Similarly, the influence of the alignment in the build chamber or on the build platform has not been considered holistically in these publications.

Similarly, the influence of the alignment in the build chamber or on the build platform has not been considered holistically in these publications. Furthermore, the influence of different toolpath parameters on the dimensional accuracy or roundness has not yet been covered by the literature. In conclusion, it is not possible to clarify from these investigations how deviations in cylindrical elements emerge and how deviations are distributed over the circumference.Experimental SetupMaterial and MachineThis chapter describes the experimental studies that are necessary to investigate the geometrical accuracy of holes and cylinders manufactured with FDM. The used model material is an ABS-M30 (natural color) with the corresponding support material SR-30 from Stratasys. ABS-M30 consists of 70-75 % polymer MABS (CAS No. 9010-94-0) and 25-30 % polymer SAN (CAS No. 9003-54-7) [21].

ABS-M30 consists of 70-75 % polymer MABS (CAS No. 9010-94-0) and 25-30 % polymer SAN (CAS No. 9003-54-7) [21]. In general, ABS (as well as modified mixtures) is an amorphous thermoplastic and therefore the shrinkage is very small (0.4 to 0.7 %) in the processing process because it has no crystalline structures [22]. A Stratasys Fortus 400mc large with a build envelope of 406 x 355 x 406 mm³ (X, Y, Z) is used to manufacture the test specimens. The layer thickness is 0.1778 mm because a T12 nozzle with a bore diameter of 0.3048 mm is used. In the data processing software Insight (version 10.4), a solid filling of the parts as well as the support style “sparse” is adjusted. The filling of the part is done at an angle of 45 ° (raster angle) and this angle changes by 90 ° from one layer to the next (δ – angle).

The filling of the part is done at an angle of 45 ° (raster angle) and this angle changes by 90 ° from one layer to the next (δ – angle). Unless otherwise specified, the toolpath parameters given in Table 1 are used.Table 1: Used toolpath and process parameters for manufacturing specimens of ABS-M30Layer thickness0.1778 mmShrink factorX, YZ1.00551.0059Contour & raster width 0.3556 mmRaster angle45 °δ - angle90 °Air gap0 mmTest SpecimensAs mentioned at the outset, the dimensional accuracy of holes and cylinders with different nominal diameters is to be determined in this paper. In order to cover a wide range of nominal diameters, these are derived from DIN EN ISO 286-1 and DIN 16742, respectively. These standards are used when values for basic tolerances are to be specified. From the listed nominal dimension classes in the standard, each maximum is selected.

These standards are used when values for basic tolerances are to be specified. From the listed nominal dimension classes in the standard, each maximum is selected. Therefore nominal diameters between 3 and 80 mm are obtained, which are used as external and internal dimensions on the test specimens (cf. Figure 2 & Figure 3). The test specimen with seven different cylinders (outer diameter) is shown in Figure 2. The cylinders with a wall thickness of 10 mm (Z-height) are arranged on a 1 mm thick base plate. The chosen arrangement results from the measurement and2759manufacturing specifications, so that they can be measured with a CMM and manufactured on a FDM system.t = 10Ø 6t = 1Ø 80Ø 50Ø 30Ø 18Ø 10 Ø 3Figure 2: Specimen with seven cylinders with nominal diameters from 3 to 80 mmSeven different inner diameters (bores) are measured on the test specimen shown in Figure 3. The arrangement of the holes is identical to the previously described test specimen (cf.

The arrangement of the holes is identical to the previously described test specimen (cf. Figure 2) with the specification that exactly 10 mm of material is arranged around and between the holes. This ensures that a constant volume is present around the bore and no influence of the external volume is included in the examination. Due to the fact that the total height of the specimen is 11 mm and a continuous base plate is used, the depth of the holes is 10 mm.t = 1Ø 6t = 11Ø 80Ø 50Ø 30Ø 18Ø 10 Ø 3Figure 3: Specimen with seven holes with nominal diameters from 3 to 80 mmDepending on the part alignment, different geometrical properties result, so that this characteristic of the FDM process is taken into account in these investigations [8-10]. Therefore, test specimens are manufactured along both the X- and Y-axes and upright in the Z-direction (see Figure 4).

Therefore, test specimens are manufactured along both the X- and Y-axes and upright in the Z-direction (see Figure 4). The test specimens in the Z-alignment require support material on both the holes and cylinders, since overhangs must be supported at an angle > 45°. In the case of the flat-lying2760specimens (X- and Y-alignment), support material is only required underneath in order to be able to remove the specimens from the build sheet. For sufficient statistics, all specimens are manufactured three times with the same batch of material. Each build job only contains specimens for one alignment, and a sacrificial tower (function of the software Control Center) is added to increase the part quality. After manufacturing, the support material is removed by dissolving in an alkaline bath so that no damage can be caused by mechanical removal.

After manufacturing, the support material is removed by dissolving in an alkaline bath so that no damage can be caused by mechanical removal. Subsequently, the parts are dried and conditioned for 88 h in standard climate according to DIN EN ISO 291.ZZ-alignmentYXX-alignmentY-alignmentFigure 4: Representation of the three considered alignments: X-, Y- and Z-alignmentMeasuring MethodThe measuring process consists, among other things, of the parameter measuring equipment, the measuring method and the examiner. Therefore, the selection of the measuring instrument and method has a decisive influence on the result. The result thus consists of a functional relationship between measured variables and influencing variables [23]. A CMM of type Altera 8.7.6 from Nikon is used to obtain the most accurate measurement possible. The measuring head of the CMM is a PH10T Plus from Renishaw with a probe of ruby (diameter 3 mm, length 21 mm).

The measuring head of the CMM is a PH10T Plus from Renishaw with a probe of ruby (diameter 3 mm, length 21 mm). The software Camio 8.2 is used to align the specimens and to program the execution of the measurement. An important point in the measurement is the choice of measuring points. In addition to the point-by-point detection of a few measuring points for a cylindrical element, this is also possible by means of a tactile scan. In this case, the probe remains permanently on the surface of the measured element and thereby records a large number of measured values. This measurement method is used in this paper in order to provide an accurate measurement as well as a statement about the occurrence of the deviations over the circumference. The measurement on the test specimen is 1.5 mm below the part surface (top layer).

The measurement on the test specimen is 1.5 mm below the part surface (top layer). For measuring by means of CMM, the specimens are aligned with their longest, continuous side at an aluminum bar and fixed with plastic clamping elements without any deformation. The CMM software Camio allows different measurement algorithms to be used in order to output absolute values. However, in this paper only the least squares method is applied. The diameter (dimensional deviation) and roundness (form deviation) are determined from the measurements (see Figure 5). The roundness is the radial distance between two boundary circles with a common center. Using the example of a cylinder of solid material, the inner circle is in the2761material and is described by the radius (r min) with the smallest distance to the center.

Using the example of a cylinder of solid material, the inner circle is in the2761material and is described by the radius (r min) with the smallest distance to the center. The outer boundary circle is completely outside and thus describes the maximum radius (rmax) [12,24].Since the 3 mm hole cannot be measured with the probe described above (diameter 3 mm), the optical measurement system VR-3100 from Keyence is used. First, an 80x magnified image is taken and a measurement circle is manually fitted. The diameter of the hole can be determined using this measurement circle.a)b)Diameter (least squares method)Outer CircleRoundness frmaxReal Measured CircleInner CirclerminReal Measured CircleFigure 5: Measurement of a) diameter by the least squares method and b) roundnessResults & DiscussionPart AlignmentThe influence of the previously described part alignments on the resulting dimensional deviations is analyzed in Figure 6.

In the left part of Figure 6 a), the deviations from the nominal dimension of holes for the alignments X, Y, Z are shown. The dotted lines represent a linear trend and the error bars indicate the standard deviation in the positive and negative direction. For internal diameters of up to 18 mm, an undersized measurement value can be recognized for all alignments. As the diameters increase, the measured values increase as well. This is also confirmed by rising trend lines. The differences for X- and Y-alignment are very small from a nominal length of 18 mm. Before this nominal diameter the differences are marginal. The dimensional deviations for all diameters of the Z-aligned specimens are initially over the flat- lying specimens (Ø ≤ 10 mm). From a nominal diameter of 30 mm, the deviations between Z and Y-alignment are only slightly different from each other. In the right part of Figure 6, the dimensional deviations over nominal diameters of the considered cylinders are plotted.

In the right part of Figure 6, the dimensional deviations over nominal diameters of the considered cylinders are plotted. In contrast to the inner diameters, almost exclusively negative deviations are measured (except the Z-aligned specimen with an 80 mm nominal diameter). It is observed that the deviations for X- and Y-alignment are identical. In the Z-alignment, only very slight deviations from the nominal diameter occur with vertically built cylinders. Only the very small nominal diameters result in an undersize (Ø ≤ 10 mm). The trend lines document a general increase in the measured values.

This is the only way to make a holistic statement about the dimensional accuracy of holes and cylinders. Figure 7 shows the results for the roundness of holes (left – a) and cylinders (right – b) over the respective nominal diameter. For holes that are manufactured in the X- and Y-alignments, no appreciable differences are obtained up to a nominal length of 30 mm. The average roundness is approx. 0.1 mm (see Figure 7). With increasing nominal diameter, a very slight increase in roundness occurs up to 0.17 mm for specimens with the largest examined diameter of 80 mm (Y-alignment). In contrast, the Z- aligned specimens have higher form deviations for all nominal diameters. Since the holes in this orientation are imaged through layers and not by a contouring filament deposition, higher deviations occur up to max. 0.28 mm (Ø 80 mm). A positive increase in the form deviation with increasing nominal diameter is apparent.

0.28 mm (Ø 80 mm). A positive increase in the form deviation with increasing nominal diameter is apparent. For cylinders manufactured in the XY-build plane, the resulting roundness is nearly constant at 0.15 mm (cf. Figure 7). Only the very small outer diameters of 3 and 6 mm deviate from this trend and have lower or higher roundness. Thus, no differences between the X- and Y-alignments can be identified. The roundness of cylinders manufactured in the Z-alignment increases with increasing nominal diameter. At first, the roundness is about 0.11 mm and continuously increases to a maximum value of 0.32 mm. These large deviations can also be explained by the layer-by-layer manufacturing process. Furthermore, support material is necessary for manufacturing the cylinders in the Z-alignment, which leads to certain deviations.

Furthermore, support material is necessary for manufacturing the cylinders in the Z-alignment, which leads to certain deviations. Depending on the quality of the calibration between model and support material, larger or smaller deviations can arise.276380X-AlignmentY-AlignmentZ-Alignment0.40 0,400.35 0,35a) Holesb) Cylindersm m0.30 0,30/s s e n d n u o R0.25 0,250.20 0,200.15 0,150.10 0,100.05 0,050,00 0.00010 20 30 40 50 60 70 80010203040506070Nominal Diameter / mmFigure 7: Influence of the part alignment on the resulting roundness: a) holes (inner diameter), b) cylinders (outer diameter)Part VolumeIn order to investigate how surrounding material affects the internal diameter of a hole, a slightly modified specimen was developed (see Figure 8). The specimen is similar to the one shown in Figure 2. However, the specimen shown in Figure 8 consists of six hollow cylinders with an internal diameter of between 6 and 80 mm.

The specimen is similar to the one shown in Figure 2. However, the specimen shown in Figure 8 consists of six hollow cylinders with an internal diameter of between 6 and 80 mm. The height of the cylinders is 10 mm and the wall thickness around the cylinders is 2 mm. This specimen is manufactured three times per alignment and measured as described above.Figure 8: Specimen with hollow cylinders with inner diameters from 6 to 80 mm (CAD model)The results of the measured diameters of holes and hollow cylinders are shown in Figure 9. For the nominal diameters 6 and 10 mm, no influence of surrounding material can be seen in the resulting deviations of specimens aligned along the X-axis. From a nominal diameter of 18 mm, a clear influence of the volume on the inner diameter is conspicuous. This is shown, inter alia, by the large slope of the linear trend line. Similarly, no effect can be seen for Y-aligned specimens with inner diameters up to 18 mm.

This is shown, inter alia, by the large slope of the linear trend line. Similarly, no effect can be seen for Y-aligned specimens with inner diameters up to 18 mm. For larger nominal diameters, the dimensional deviations of holes and hollow cylinders diverge, with the exception of the nominal diameter 50 mm. For the specimens manufactured in the Z-alignment, an influence is recognizable as well. As already shown in Figure 6, the dimensional deviations increase with increasing nominal diameter. On the other hand, the hollow cylinders have very constant deviations of about 0.03 mm over the entire diameter range. Thus it can be stated here that the internal diameter of hollow cylinders is always276480much smaller.

Thus it can be stated here that the internal diameter of hollow cylinders is always276480much smaller. However, for each nominal length it is necessary to prove independently where better dimensional accuracy exists.0.20 0,20X (Hole) Y (Hole)Z (Hole) X (Hollow) Y (Hollow)Z (Hollow)m m0.15 0,15/ n o i t a i v e D0.10 0,100.05 0,050.00 0,00l a n o i s n e m Di0,05 -0.050,10 -0.100,15 -0.150,20 -0.2001020304050607080Nominal Diameter / mmFigure 9: Influence of the surrounding part volume on the resulting dimensional deviationsThe surrounding volume has no influence on the roundness for specimens which are manufactured in the X- and Y-alignment. For the Z-alignment, it depends on the nominal diameter which specimen has a lower form deviation, so that no clear statement can be made.Air GapThe toolpath parameter “air gap” specifies the distance between two deposited filaments.

By means of a positive value, a lightweight structure can be created since a lower filling of the part is achieved. With a negative air gap it is possible to increase the mechanical properties of FDM parts [7]. As the distance between two filaments becomes negative, there is a slight overlap of the filaments and thus increased filling of the part is realized. If too large a negative value is chosen, an overfilling of the part may occur. This impairs the dimensional accuracy and surface finish. In these investigations, the air gap was set at -0.017 mm, which corresponds to 5% of the filament width of 0.3556 mm. The purpose of this study is to determine whether the negative air gap has an influence on the dimensional or form deviations of cylindrical elements. The dimensional deviations are shown in Figure 10 for holes (left – a) and cylinders (right – b) over the considered nominal diameters.

The dimensional deviations are shown in Figure 10 for holes (left – a) and cylinders (right – b) over the considered nominal diameters. The slightly negative air gap has an effect on X-aligned test specimens with holes when the diameter is less than 30 mm. In this range, the mean values of the deviations are slightly larger than for the standard parameters (air gap = 0 mm). For holes with a diameter larger than 30 mm, no difference can be detected. However, test specimens with holes, manufactured in the Y-alignment and with a negative air gap, show slightly higher values over all nominal diameters. One possible reason for these observations could be a larger shrinkage due to the increased filling of the part.

Although the trend lines are slightly offset from one another, this is due to the outlier at a diameter of 3 mm (see Figure 10). Figure 11 shows the views of the toolpath as well as microscopy images. The dimensional deviation in a cylinder with a diameter of 3 mm results from overfilling. In comparison to standard parameters, one filament is placed more in the cylinder (cf. Figure 11 – a & c). In the red circle of Figure 11 – d), it is noted that a form error occurs at the position of the weld line which leads to the observed deviation. For specimens with a larger nominal diameter, this effect is not critical, since the total part area is larger.

For specimens with a larger nominal diameter, this effect is not critical, since the total part area is larger. In the design process, it must be ensured that dimensional and form deviations can occur in delicate part areas if too large a negative air gap is selected.a)b)c)d)1 mm1 mmFigure 11: Influence of the air gap on a cylinder with diameter 3 mm: a) Insight toolpath - standard parameter, b) microscopy image – air gap 0 mm, c) Insight toolpath – air gap 0 mm, d) microscopy image – negative air gap.The influence of a negative air gap on the roundness of the hole and cylinder is analyzed in the following diagram (cf. Figure 12 - a). In the case of X-aligned manufactured holes, the negative air gap has no influence on the resulting roundness for diameters up to 30 mm. For larger nominal diameters, there is a tendency for the negative air gap to result in slightly higher form deviations.

For larger nominal diameters, there is a tendency for the negative air gap to result in slightly higher form deviations. For specimens that have been aligned and manufactured along the Y-axis, there276680is no difference in the roundness of the inner diameter due to a changed air gap. In the right-hand part of Figure 12, the form deviation for the cylinder specimens in the alignments X and Y is examined. The highest form deviation is recorded for a nominal diameter of 3 mm in the X- alignment and amounts to 0.23 mm. As the nominal diameter increases, the roundness improves and remains at a constant level of approx. 0.12 mm (Ø 50 and 80 mm). There is no outlier for the Y-alignment, and no differences between negative and zero air gap are visible. Summarizing, it can be stated here that the air gap has no or only a small effect on the roundness.

Summarizing, it can be stated here that the air gap has no or only a small effect on the roundness. If a tiny influence is noticeable, then only at very small nominal diameters.X (Air Gap)X (Standard)Y (Air Gap)Y (Standard)0.40 0,400.35 0,35a) Holesb) Cylindersm m0.30 0,30/s s e n d n u o R0.25 0,250.20 0,200.15 0,150.10 0,100.05 0,050,00 0.00010 20 30 40 50 60 70 80010203040506070Nominal Diameter / mmFigure 12: Influence of the toolpath parameter “air gap” on the resulting roundness: a) holes (inner diameter), b) cylinders (outer diameter)Shrink FactorIn previous publications it was shown that the shrink factor (SF) has plenty of potential to reduce the dimensional deviations and improve the accuracy [9]. The diagrams in Figure 6 and Figure 7 showed that an undersize occurs for almost all alignments with a nominal diameter smaller than 30 mm.

The diagrams in Figure 6 and Figure 7 showed that an undersize occurs for almost all alignments with a nominal diameter smaller than 30 mm. Since holes with a diameter of up to 30 mm are used on a large scale, the shrink factor is increased by an amount of + 0.001. For the flat build-up directions X and Y, the SF was changed by the same value (SFx = SFy = 1.0065) and the Z-shrink factor was increased for the Z-alignment (SFz = 1.0069). In Figure 13, the occurring dimensional deviations for specimens with increased SF (SF_X+Y+ or SF_Z+) are compared with default shrink factors. In the left section, the dimensional deviations for the specimens with holes are shown and in the right section, the deviations for the specimens with cylinders. For holes, no influence can be detected, irrespective of the alignment on the build platform, if the nominal diameter is 6 or 10 mm. As the nominal diameter increases (≥ 18 mm), the difference between raised and default shrink factors increases as well.

As the nominal diameter increases (≥ 18 mm), the difference between raised and default shrink factors increases as well. This observation was expected as the increased shrink factor leads to a larger scale of the CAD model. Therefore, this increase is also expected to result in higher measured values. It is also direction-dependent whether an improvement of the accuracy can be achieved or not. The dimensional deviations over the nominal diameter of the measured cylinders are shown in the right part of Figure 13. For cylinders with a diameter of 3 mm (X- and Y-alignments), an increase in the measured value and thus an improvement in dimensional accuracy can be observed. Analogous to the results of the holes, there is likewise no influence of276780a higher shrink factor on specimens with a diameter of 6 and 10 mm. This applies both to the X- and Y-alignments as well as to the upright manufactured specimen (Z).

This applies both to the X- and Y-alignments as well as to the upright manufactured specimen (Z). In all three alignments, the measured diameters (Ø ≥ 18 mm) increase significantly with increased shrink factor. In the X- and Y-alignments, an improvement of the dimensional accuracy of cylinders is possible for nominal diameters up to 80 mm. The Z-alignment already has a good dimensional accuracy with standard shrink factors, so there is no improvement at this point. The respective uniform increase of the shrink factor has no influence on the roundness of the specimens.

Depending on the part, a different shrink factor can be used in the X-alignment than in the Y-alignment. However, if there is a cylindrical element in the part, the non-uniform adjustment of the shrink factor may lead to dimensional and form deviations. The purpose of the following investigation is to determine how an uneven increase of the shrink factor affects the dimensional accuracy as well as the roundness. Figure 14 shows the dimensional deviations for specimens that have been manufactured and measured with unevenly and evenly adapted shrink factors. SF_Y+ or SF_X+ stands for an evenly increased value to SFx = SF y = 1.0065. In this figure, only the X-alignment is shown because the results for the Y-alignment have the same tendencies. For the element hole, identical dimensional deviations occur for unevenly (SF_Y+, SF_X+) and evenly (SF_X+Y+) increased shrink factors (cf. Figure 14 a). This is also illustrated by the linear trend lines.

Figure 14 a). This is also illustrated by the linear trend lines. A different image is obtained for cylinders because the measured values for evenly increased SF (SF_X+Y+) are higher for nominal diameters ≥ 18 mm (Figure 14 b). The unevenly increased shrink factors only lead to a small increase in the measured values compared to the standard factors (compare green and blue trend line versus purple trend line). Again, this is only observable for higher nominal diameters. A slight increase in the measured values can be explained by the increased scaling. However, this effect is dependent on the considered element.2768In the next step, the effects of the unevenly increased shrink factors on the roundness are analyzed (see Figure 15). Again, only the results for X-aligned specimens are shown, since the results are similar to specimens in the Y-alignment.

Again, only the results for X-aligned specimens are shown, since the results are similar to specimens in the Y-alignment. The linear trend lines (dotted lines in the left diagram of Figure 15) indicate a slight increase of form deviations with increasing nominal diameter of the hole due to unevenly increased shrink factors. However, the increase is very slight in places and is often within the standard deviation of the measured values from the unchanged shrink factors.

However, the increase is very slight in places and is often within the standard deviation of the measured values from the unchanged shrink factors. For the manufactured cylinders (right diagram in Figure 15), no influence in the X- and Y-alignments is visible due to unevenly increased shrink factors on the roundness.m m0.20 0,200,15 0.15SF_Y+a) HolesSF_X+SF_X+Y+SF_defaultb) Cylinders/n o i t a i v e D0,10 0.100.05 0,050.00 0,00l a n o i s n e m Di0,05 -0.050,10 -0.100,15 -0.150,20 -0.200102030405060 80 Nominal Diameter / mm70010203040506070Figure 14: Influence of non-uniformly increased shrink factors (SFx & SFy) on the resulting dimensional deviations: a) holes (inner diameter), b) cylinders (outer diameter) – X-alignmentAs a result, it can be said that there are only slight, hardly significant changes of the roundness if the shrink factors are unevenly increased.

Due to the continuous recording of measuring points during the measurement by means of CMM (tactile scan), it is possible to make an accurate statement about the occurrence of the dimensional deviations for the cylindrical element. Figure 16 shows the dimensional deviations over the circumference for X-aligned holes and cylinders with a nominal diameter of 50 mm. These parts were manufactured with standard parameters (cf. Table 1). The measurement was made 1.5 mm below the upper part surface. It is striking that the highest deviations always occur at the place of the weld line (see red circles in Figure 16). Since the start and end points of the deposited filament coincide here, a certain accumulation of material is produced which leads to visible and measurable deviations of the ideal contour. Furthermore, it is noticeable that at certain points of the circumference, an undersize or oversize emerges.

Furthermore, it is noticeable that at certain points of the circumference, an undersize or oversize emerges. Taking into account the FDM toolpath (Figure 17), it is possible to find correspondences between the direction of filament deposition and occurring deviations. The manufactured specimens in the Y-alignment have identical dimensional deviations over the circumference but are each rotated through 90°.2770X-alignment - Nominal Diameter 50 mma) Hole – Inner Diameterb) Cylinder – Outer DiameterWeld LineFigure 16: Course of the deviations for a) hole / inner diameter and b) cylinder / outer diameterIn the areas in which the deposited filament connects directly to the contour (cf. Figure 17 – Red Circle 1), an oversize occurs. An explanation for this effect may be the shrinkage of the filament which leads to a material contraction due to cooling. As a result, the contour is easily pulled away from its ideal position and results in the deviation profile shown here.

As a result, the contour is easily pulled away from its ideal position and results in the deviation profile shown here. Due to the filament deposition at an angle of 45°, the filaments have little or no effect on the contour of the hole in other areas (cf. Figure 17 – Red Circle 2). As a result there is an undersize in these areas. In the case of cylinders (outer diameter – cf. Figure 16 b), it is exactly the opposite, because the connection of the filaments is realized from inside to the contour.a)b)2)X-alignmentDirection of Deposition1)Layer X Layer X+1Figure 17: Analysis of the FDM-toolpath for flat built specimens (X- and Y-alignment)Considering the filament deposition, logical deviations result over the circumference of the hole. However, it is necessary to take into account the alternating filament deposition from one layer to the next by 90° (δ - angle), whereby the deviations would actually have to be canceled.

However, it is necessary to take into account the alternating filament deposition from one layer to the next by 90° (δ - angle), whereby the deviations would actually have to be canceled. Due to the constant measuring height of 1.5 mm below the upper part surface, the measurement could only have been carried out on a single layer. To exclude this influence, the hole was measured at a depth of 3, 6 mm and a combination of both. However, as Figure 18 shows, the areas where the oversize or undersize occurred were identical. In addition, a helical scan over 6 mm of the hole surface was performed. The probe starts at a depth of 1.5 mm below the surface of the part and moves the hole contour in a spiral. As a result, the previously determined trend can also be detected, but in attenuated form (cf. Figure 18 e).

As a result, the previously determined trend can also be detected, but in attenuated form (cf. Figure 18 e). Thus the influence of the delta- angle on the resulting deviations over the circumference cannot be conclusively clarified.2771Figure 18: Deviations over the circumference of a hole (Ø 50 mm) for different measuring heightsTherefore further investigations with different filling strategies (variation of the raster and delta angle) are carried out. As shown in Figure 19, holes and cylinders with a nominal diameter of 80 mm are manufactured with different angles and measured as described above. The first value indicates the direction of the filament deposition to the X-axis. The second value is the δ- angle, which describes the change of the filament deposition from one layer to the next. A δ-angle of 0° means that there is no change in the direction of deposition. Correspondingly, 90° means an alternating deposition.

A δ-angle of 0° means that there is no change in the direction of deposition. Correspondingly, 90° means an alternating deposition. The deviations show a uniform trend that the deviations adapt to the raster angle. This results in an undersize in the direction of filament deposition, which is caused by shrinkage effects.90° / 0°90° / 90°60° / 90°45° / 90°ye l o Hxr e d n i l y CFigure 19: Influence of different raster and delta angles on the resulting deviations over the circumferenceIn addition to the manufactured specimens in the XY-plane, the deviations of Z-aligned specimens with holes and cylinders are shown in Figure 20. The deviations over the circumference are shown for a nominal diameter of 50 mm: on the left for an inner diameter (hole) and on the right for an outer diameter (cylinder). The profiles of the occurring deviations are clearly different from the specimens manufactured in the X- and Y-alignment.

The profiles of the occurring deviations are clearly different from the specimens manufactured in the X- and Y-alignment. The influence of the support material, which is used in or under the cylindrical element, is obviously recognizable by increased dimensional deviations. For holes, very large local undersize occurs below and above due to the use of support material. In contrast, a locally large oversize occurs in2772the left and right-hand part of the hole. These dimensional deviations can also be explained by shrinkage effects, since the surrounding material contracts due to cooling. Similarly, support material is required for cylinders in the Z-alignment, which is particularly visible by local oversize in the lower region (cf. Figure 20 – b).

Similarly, support material is required for cylinders in the Z-alignment, which is particularly visible by local oversize in the lower region (cf. Figure 20 – b). Orthogonally, a local undersize appears which becomes smaller as the number of layers increases.Z-alignment - Nominal Diameter 50 mma) Hole – Inner Diameterb) Cylinder – Outer DiameterzFigure 20: Course of the deviations for a) hole / inner diameter and b) cylinder / outer diameterFigure 21 illustrates how the dimensional deviations over the circumference emerge for specimens manufactured in the Z-alignment. In contrast to specimens which are manufactured in the XY-plane, there is no continuous contour filament that describes the ideal contour of the element. In this orientation, the circular contour is formed by layers that are iteratively deposited.

In this orientation, the circular contour is formed by layers that are iteratively deposited. Due to the staircase effect in additive manufacturing processes on round geometries, dimensional deviations arise, which can be reduced by choosing a smaller layer thickness.a) Holeb) CylinderzModelSupportIdeal contourFigure 21: Analysis of the structure of vertical built specimens a) hole and b) cylinderSummarizing, Figure 22 shows how dimensional deviations of different cylindrical elements occur at different alignments. However, the deviations shown here apply only when using the standard filling strategy at an angle of 45°. By means of the diagram, users of the FDM process can be taught how dimensional deviations of cylindrical elements occur.

By means of the diagram, users of the FDM process can be taught how dimensional deviations of cylindrical elements occur. Furthermore, it is important for post-processing methods such as drilling, that there is enough machining allowance given to the cylindrical elements.2773Part AlignmentxyzDeviationss e l o HInternal DimensionsExternal Dimensionsr e d ni l y CFigure 22: Schematic representation of the occurring deviationsConclusions and OutlookThe aim of these investigations was to analyze and optimize the dimensional accuracy of holes and cylinders manufactured with FDM and the material ABS-M30. For this purpose, specimens were developed which contained outer and inner diameters between 3 and 80 mm. By means of a coordinate measuring machine, it was possible to gather both dimensions and form deviations. By considering different alignments, it was possible to ascertain that different dimensional deviations occur depending on the alignment.

By considering different alignments, it was possible to ascertain that different dimensional deviations occur depending on the alignment. In addition, there are different dimensional deviations for inner (hole) and outer (cylinder) diameters. These elements have one thing in common: the measured value increases with increasing nominal diameter. However, whether there is an undersize or oversize depends on the respective nominal diameter. The roundness of cylindrical elements manufactured in the XY-plane is very constant and lower than for elements manufactured in the Z-alignment. Another part of this paper deals with the influence of different process parameters, such as air gap or shrink factor, on the resulting dimensional and form deviations. By adjusting these parameters, an improvement of the dimensional accuracy was possible.

By adjusting these parameters, an improvement of the dimensional accuracy was possible. By visualizing the local dimensional deviations over the circumference of the elements, it was possible to clarify how the direction of the filament deposition (raster angle) affects the deviating deviations. It has been proved that an undersize comes about in the direction of the filament deposition. Reasons for this behavior are shrinkage effects, which lead to a contraction of the polymer filaments. In addition, the use of support material causes dimensional deviations on cylindrical elements manufactured in the Z-alignment. Finally, it could not be conclusively clarified why an alternating filament deposition (δ-angle of 90°) does not lead to a removal of the illustrated dimensional deviations. In further investigations, it will be necessary to examine how dimensional and form deviations in FDM parts appear, which are not aligned along the axes of the FDM process.

In further investigations, it will be necessary to examine how dimensional and form deviations in FDM parts appear, which are not aligned along the axes of the FDM process. It is also necessary to check whether the formation of local dimensional deviation is machine-dependent or not.2774References[1] Wohlers, T.: Wohlers Report 2015. Annual Worldwide Progress Report. 2015 [2] Acatech – National Academy of Science and Engineering, German National Academy of Sciences Leopoldina, The Union of the German Academies of Sciences and Humanities.: Additive Fertigung. Stellungnahme, 2016[3] VDI Standard 3405: Additive manufacturing processes, rapid manufacturing – Basics,definitions, processes. 2014[4] Gibson, I., Rosen, D., Stucker, B.: Additive Manufacturing Technologies – 3D Printing,Rapid Prototyping and Direct Digital Manufacturing. 2nd Edition, Springer, 2015[5] Fischer, F.: Thermoplastics: The strongest choice for 3D Printing.

ometrical Analysis of Simple Contours Deposited by a 3D Printing HexacopterAlexander Nettekovena, Nicholas Frankenb, Ufuk TopcubaWalker Department of Mechanical Engineering, The University of Texas at Austin, 204 E. Dean Keeton Street, Austin, 78712, TX, USA bDepartment of Aerospace Engineering and Engineering Mechanics, The University of Texas at Austin, 2617 Wichita Street, Austin, 78712, TX, USAAbstractCurrent limitations in vertical and horizontal mobility for ground robots in 3D printing of medium to large-scale objects have recently led to the development of a 3D printing hexacopter testbed at the University of Texas at Austin. This testbed can ﬂy to a desired location and deposit polylac- tic acid on ﬂat surfaces. A previous study has shown the feasibility of this approach but has not yet quantiﬁed the testbed’s printing capabilities. In this paper, we quantify the printing capabil- ities.

A previous study has shown the feasibility of this approach but has not yet quantiﬁed the testbed’s printing capabilities. In this paper, we quantify the printing capabil- ities. We print square contours of diﬀerent sizes and quantify the printed results based on their geometric dimensions. We also quantify the testbed’s trajectory tracking to assess the testbed’s positioning accuracy during printing. In quantifying the testbed, we lay the groundwork for using aerial robots in printing applications of medium to large-scale objects, such as concrete printing.1. IntroductionGround robots are increasingly used in 3D printing of medium to large-scale objects [1, 2, 3, 4, 5]. Pairing a robot’s versatility with 3D printing signiﬁcantly increases the process’s ﬂexibility and an engineer’s design freedom to manufacture arbitrarily complex objects [5].

Pairing a robot’s versatility with 3D printing signiﬁcantly increases the process’s ﬂexibility and an engineer’s design freedom to manufacture arbitrarily complex objects [5]. However, a lack of vertical mobility and sometimes horizontal mobility still prevents these 3D printing processes from maximizing their potential [2, 5, 6]. For example, the limitations in mobility usually require the ground robots to be larger than the to-be-printed objects [6]. This requirement prevents these systems from scaling to arbitrarily large objects. Another limitation is a ground robot’s dependency on the terrain and the surroundings. For example, a ground robot cannot print a large cylindrical structure from its center without being trapped inside.To overcome these limitations, we recently developed a 3D printing hexacopter testbed at the University of Texas at Austin (see Figure 1(a)) [7].

This testbed is a small-scale proof of concept for combining 3D printing technologies with unmanned aerial vehicles (UAV). The testbed uses fused deposition modeling to 3D print during ﬂight. The advantage of combining 3D printing with UAV technology is the added mobility in the vertical direction and the lower dependency on the terrain and the surroundings. We previously demonstrated the feasibility of this approach by successfully printing diﬀerent contours of polylactic acid (PLA) on a ﬂat surface during ﬂight. However, these tests were primarily qualitative.In this paper, we analyze the system’s printing capabilities quantitatively. We perform mul- tiple print trials and measure the geometric dimensions of the print results.

We perform mul- tiple print trials and measure the geometric dimensions of the print results. Since 3D printing applications such as concrete printing require 3D printing at speciﬁc locations, we also analyze(a) 3D Printing Hexacopter Testbed(b) Top view schematicFigure 1: The 3D printing hexacopter is depositing polylactic acid on a ﬂat surface according to the desired trajectory on the right. A video of previous experiments with the testbed can be found at https://youtu.be/tEooDpE2TyE.the testbed’s positioning accuracy. In performing these quantiﬁcations, we lay the groundwork for further development of this novel process and enable benchmarking with other 3D printing processes in future research.The remainder of this paper is structured as follows. In Section II, we introduce the methods for quantifying the current printing capabilities of the testbed. In Section III and IV, we present our results from multiple print trials and discuss their relevance.

In Section III and IV, we present our results from multiple print trials and discuss their relevance. In Section V, we provide our conclusions and discuss the next steps.2. MethodsTo quantify the printing performance of the testbed, we print three contours of diﬀerent sizes on a ﬂat surface (see Figure 1). Note that we focus on single-layer geometries in this paper since multi-layer printing has not been tested yet. The desired edge lengths of the three square contours are 60 mm, 100 mm, and 150 mm. We then use two diﬀerent methods to analyze the testbed’s printing performance. The ﬁrst method focuses on the geometric dimensions of the printed contours, while the second method focuses on the testbed’s positioning accuracy.For the ﬁrst method, we measure the edge lengths of each printed square contour with a digi- tal caliper and then compare the measured edge lengths to the desired edge lengths. The caliper’s tolerance is 0.03 mm.

The caliper’s tolerance is 0.03 mm. For the second method, we record the testbed’s position throughout the print trials and compare the measured positions with the positions given by the desired trajec- tories. In comparing the positions, we determine how accurate the positioning of the testbed is. This analysis is crucial since large-scale printing applications, such as concrete printing, often require printing at a speciﬁc print location. To measure the testbed’s position, we use a motion capture system with submillimeter position accuracy.3. ResultsWe present the successful print results of the square contours in Figure 2. Each square is marked with the corresponding edge measurements in millimeters. As indicated by the measure- ments, we ﬁnd that all but one edge are larger than the desired lengths.

Each square is marked with the corresponding edge measurements in millimeters. As indicated by the measure- ments, we ﬁnd that all but one edge are larger than the desired lengths. We calculate the root mean square error (RMSE) for the 60 mm × 60 mm, 100 mm × 100 mm, and 150 mm × 150 mm2(a) 60 mm × 60 mm square contour with an RMSE of 12.17 mm.(b) 100 mm × 100 mm square con- tour with an RMSE of 11.49 mm(c) 150 mm × 150 mm square con- tour with an RMSE of 9.12 mmFigure 2: Square contours printed by the 3D printing hexacopter. A 30-centimeter ruler is placed alongside the squares for comparison. The marked measurements of the edge lengths are in millimeters.squares to be 12.17 mm, 11.49 mm, and 9.12 mm, respectively. The total RMSE for all edges is 10.8 mm.The positioning accuracy of the testbed has a root mean square error of 4.3 mm. This posi- tioning error is considerably lower than the deviations in the edge lengths.

This posi- tioning error is considerably lower than the deviations in the edge lengths. However, the max- imum position error of the testbed, averaged over the three print trials, is 11.2 mm. From our experiments, we ﬁnd that the largest position errors, including the maximum position error for each trial, occur at the corners of the square contours and are caused by the testbed overshooting the desired corner setpoints during the print trials (similar to the exemplary traversed trajectory in Figure 1(b)).4. DiscussionWe ﬁnd that the testbed achieves approximately four-millimeter positioning accuracy during printing. This accuracy is between double and triple the reported positioning accuracy for some commercial concrete printing machines that use large-scale gantry or robotic systems. These machines can have positioning accuracies of one to two millimeters [8].

These machines can have positioning accuracies of one to two millimeters [8]. However, the current testbed setup is using an idealized laboratory environment, which reduces outside disturbances, such as wind gusts, and relies on submillimeter position estimation from a motion capture system for the ﬂight controllers. To achieve similar results in outdoor environments, future research needs to investigate the robustness of the testbed outside of the laboratory setting.When considering the geometric dimensions of the printed contours, the testbed’s tracking performance does not translate to similar accuracy in the edge lengths. The discrepancy between accuracy in positioning and geometric dimensions is largely due to the testbed overshooting at the corners of the square contours.

The discrepancy between accuracy in positioning and geometric dimensions is largely due to the testbed overshooting at the corners of the square contours. Since the testbed uses PID controllers for the positioning, complex interactions between the testbed and the build surface, such as ground eﬀects and the stick-slip phenomenon, pose signiﬁcant challenges for these controllers. To achieve millime- ter accuracy in the edge lengths, the testbed requires advanced controllers that can model the testbed’s dynamics and account for the testbed’s interactions with the build surface.3An additional observation we make is that the edge length RMSE for each square contour does not grow with the size of the square contours. In fact, the RMSE slightly decreases. This characteristic may suggest that the error in the edge lengths is independent of the contour size and can be scaled to arbitrarily large dimensions.

In fact, the RMSE slightly decreases. This characteristic may suggest that the error in the edge lengths is independent of the contour size and can be scaled to arbitrarily large dimensions. However, the number of print trials is insuﬃcient to make any deﬁnitive claims.Overall, the testbed can achieve approximately four-millimeter positioning accuracy and can print contours with approximately one-centimeter tolerances. Future research will aim to im- prove the single-layer printing capabilities and also test multi-layer printing.5. ConclusionA recently developed 3D printing hexacopter testbed can deposit polylactic acid on ﬂat sur- faces during ﬂight. To quantify the testbed’s printing capabilities, we printed square contours of diﬀerent sizes and analyzed the printed results based on their geometric dimensions. We also an- alyzed the testbed’s positioning accuracy during printing.

We also an- alyzed the testbed’s positioning accuracy during printing. We ﬁnd that the testbed’s positioning accuracy has a tolerance of a few millimeters with a root mean square error of 4.3 mm. However, we ﬁnd that the printed edges are approximately 1 cm larger than the desired edge lengths. These length deviations are mainly due to the testbed overshooting at the corners of the square contours. In future research, we will improve the positioning accuracy and address the overshoot problems by developing advanced control algorithms.AcknowledgmentsThis material is based upon work supported by the National Science Foundation under GrantNo. 1646522 and 1652113.References[1] K. Wong, A. Hernandez, A Review of Additive Manufacturing, ISRN Mechanical Engineering 2012 (08 2012).doi:10.5402/2012/208760.[2] F. Bos, R. Wolfs, Z. Ahmed, T.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceGeometrical Analysis of Simple Contours Deposited by a 3D Printing HexacopterAlexander Nettekovena, Nicholas Frankenb, Ufuk TopcubaWalker Department of Mechanical Engineering, The University of Texas at Austin, 204 E. Dean Keeton Street, Austin, 78712, TX, USA bDepartment of Aerospace Engineering and Engineering Mechanics, The University of Texas at Austin, 2617 Wichita Street, Austin, 78712, TX, USAAbstractCurrent limitations in vertical and horizontal mobility for ground robots in 3D printing of medium to large-scale objects have recently led to the development of a 3D printing hexacopter testbed at the University of Texas at Austin. This testbed can ﬂy to a desired location and deposit polylac- tic acid on ﬂat surfaces.

This testbed can ﬂy to a desired location and deposit polylac- tic acid on ﬂat surfaces. A previous study has shown the feasibility of this approach but has not yet quantiﬁed the testbed’s printing capabilities. In this paper, we quantify the printing capabil- ities. We print square contours of diﬀerent sizes and quantify the printed results based on their geometric dimensions. We also quantify the testbed’s trajectory tracking to assess the testbed’s positioning accuracy during printing. In quantifying the testbed, we lay the groundwork for using aerial robots in printing applications of medium to large-scale objects, such as concrete printing.1. IntroductionGround robots are increasingly used in 3D printing of medium to large-scale objects [1, 2, 3, 4, 5]. Pairing a robot’s versatility with 3D printing signiﬁcantly increases the process’s ﬂexibility and an engineer’s design freedom to manufacture arbitrarily complex objects [5].

We perform mul- tiple print trials and measure the geometric dimensions of the print results. Since 3D printing applications such as concrete printing require 3D printing at speciﬁc locations, we also analyze1549(a) 3D Printing Hexacopter Testbed(b) Top view schematicFigure 1: The 3D printing hexacopter is depositing polylactic acid on a ﬂat surface according to the desired trajectory on the right. A video of previous experiments with the testbed can be found at https://youtu.be/tEooDpE2TyE.the testbed’s positioning accuracy. In performing these quantiﬁcations, we lay the groundwork for further development of this novel process and enable benchmarking with other 3D printing processes in future research.The remainder of this paper is structured as follows. In Section II, we introduce the methods for quantifying the current printing capabilities of the testbed. In Section III and IV, we present our results from multiple print trials and discuss their relevance.

The caliper’s tolerance is 0.03 mm. For the second method, we record the testbed’s position throughout the print trials and compare the measured positions with the positions given by the desired trajec- tories. In comparing the positions, we determine how accurate the positioning of the testbed is. This analysis is crucial since large-scale printing applications, such as concrete printing, often require printing at a speciﬁc print location. To measure the testbed’s position, we use a motion capture system with submillimeter position accuracy.3. ResultsWe present the successful print results of the square contours in Figure 2. Each square is marked with the corresponding edge measurements in millimeters. As indicated by the measure- ments, we ﬁnd that all but one edge are larger than the desired l engths.

Each square is marked with the corresponding edge measurements in millimeters. As indicated by the measure- ments, we ﬁnd that all but one edge are larger than the desired l engths. We calculate the root mean square error (RMSE) for the 60 mm × 60 mm, 100 mm × 100 mm, and 150 mm × 150 mm1550(a) 60 mm × 60 mm square contour with an RMSE of 12.17 mm.(b) 100 mm × 100 mm square con- tour with an RMSE of 11.49 mm(c) 150 mm × 150 mm square con- tour with an RMSE of 9.12 mmFigure 2: Square contours printed by the 3D printing hexacopter. A 30-centimeter ruler is placed alongside the squares for comparison. The marked measurements of the edge lengths are in millimeters.squares to be 12.17 mm, 11.49 mm, and 9.12 mm, respectively. The total RMSE for all edges is 10.8 mm.The positioning accuracy of the testbed has a root mean square error of 4.3 mm. This posi- tioning error is considerably lower than the deviations in the edge lengths.

These machines can have positioning accuracies of one to two millimeters [8]. However, the current testbed setup is using an idealized laboratory environment, which reduces outside disturbances, such as wind gusts, and relies on submillimeter position estimation from a motion capture system for the ﬂight c ontrollers. To a chieve s imilar r esults i n o utdoor e nvironments, f uture research needs to investigate the robustness of the testbed outside of the laboratory setting.When considering the geometric dimensions of the printed contours, the testbed’s tracking performance does not translate to similar accuracy in the edge lengths. The discrepancy between accuracy in positioning and geometric dimensions is largely due to the testbed overshooting at the corners of the square contours.

The discrepancy between accuracy in positioning and geometric dimensions is largely due to the testbed overshooting at the corners of the square contours. Since the testbed uses PID controllers for the positioning, complex interactions between the testbed and the build surface, such as ground eﬀects and the stick-slip phenomenon, pose signiﬁcant c hallenges f or t hese c ontrollers. To a chieve millime- ter accuracy in the edge lengths, the testbed requires advanced controllers that can model the testbed’s dynamics and account for the testbed’s interactions with the build surface.1551An additional observation we make is that the edge length RMSE for each square contour does not grow with the size of the square contours. In fact, the RMSE slightly decreases. This characteristic may suggest that the error in the edge lengths is independent of the contour size and can be scaled to arbitrarily large dimensions.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperGeometrical Deviations In Additive Manufacturing – Influences On The Manufacturing AccuracyT. Lieneke*,†, S. Lammers*,†, D. Zimmer*,†Chair of Design and Drive Technology, Paderborn University, 33098 Paderborn, Germany †Direct Manufacturing Research Center / Paderborn Institute for Additive Fabrication, Paderborn University, 33098 Paderborn, GermanyAbstractThe advantages of Additive Manufacturing (AM) highlight the capability to become an inherent part within the product development process. However, process specific challenges harm its further currency for industrial applications, for instance the high geometrical deviations. Different process factors influence the manufacturing accuracy and lead to large dimensional, form and positional deviations.

Different process factors influence the manufacturing accuracy and lead to large dimensional, form and positional deviations. Published research relative to deviations is difficult to compare, because it is based on several specimens that were manufactured with different processes, materials and machine settings. This fact emphasizes that reliable tolerance values for AM are hard to define in standards. Within this investigation, a universally applicable method was developed to examine geometrical deviations for AM processes. The main aim is the derivation of achievable tolerance values considering important influencing factors. Furthermore, due to the locally varying surface roughness of additively manufactured parts several tactile measurements were compared.IntroductionAdditive Manufacturing (AM) has been known since the 1980s.

The development of the AM processes illustrates that the technology has high technical and economical potentials that become more and more visible due to the increased market volume in the last decades. Apart from the current challenges the COVID-19 pandemic can cause an expected decrease between 10% and 39% in 2021 [1]. However, ongoing machine, material and research developments will enable a further growth in the market. AM processes produce components and assemblies by joining single volumes, usually expressed as layers.

AM processes produce components and assemblies by joining single volumes, usually expressed as layers. They do not require molding tools and the layer-by-layer manufacturing breaks down complex three-dimensional challenges into two-dimensional, simple manufacturing situations, allowing the processes to offer enormous freedom in the areas of material science and design [2, 3].Although the technical relevance of the processes has grown tremendously in the recent years, several process-specific challenges inhibit the increasing distribution and use of the processes for serial applications outside the leading industry sectors, such as aerospace, medical or automotive [4]. One essential drawback is the high surface roughness of additively manufactured components, which arises, for example, as a result of the stair-stepping effect.

One essential drawback is the high surface roughness of additively manufactured components, which arises, for example, as a result of the stair-stepping effect. In addition to the surface roughness, the components often exhibit higher dimensional, form and positional deviations compared to components that are manufactured with conventional processes. Due to the lack of knowledge regarding the achievable manufacturing accuracy, additively manufactured components are often provided with large oversize, which causes an increase of costs in the manufacturing process and in post-processing to meet the geometrical requirements of the components.437Publications on the subject of geometrical accuracy often show a qualitative evaluation of existing processes using standard benchmark parts and do not include detailed quantification of achievable tolerance values.

The main target of these investigations are the repeatability, manufacturing speed or the manufacturability of minimal geometries. Existing investigations are difficult to compare, since they refer to different machines, materials, process parameters, test specimen and measuring methods. Particularly in the area of the measurement method, specially adapted extraction strategies must be used to record deviations realistically. One reason for individual numbers and positions of measuring points is the high and inhomogeneous surface roughness on a single component.Overall, the manufacturing accuracy in AM is insufficiently researched, although many different benchmark artifacts for the evaluation of geometric accuracy for additive manufacturing processes have been investigated. REBAIOLI and FASSI provide an intensive overview of the existing investigations in this area [34].

REBAIOLI and FASSI provide an intensive overview of the existing investigations in this area [34]. Nevertheless, many of these investigations neither consider Geometric Dimensioning & Tolerancing (GD&T) characteristics [35] nor a systematic selection of geometric features as well as the quantification of geometric tolerances [36, 37]. Furthermore, there are special requirements for additively manufactured components that have not yet been fully described in the state of the art or standards, for example the tolerancing of complex free-form surfaces, lattice structures or support structures [38]. In addition, measurement methods must be adapted to the AM specific challenges, such as characteristic surface texture with high surface roughness [38]. Furthermore, causal relationships between the occurring deviations and the process-specific environment are often missing, since a multitude of influencing factors must be considered.

Furthermore, causal relationships between the occurring deviations and the process-specific environment are often missing, since a multitude of influencing factors must be considered. A large number of investigations are focused on optimizing process parameters and machine components for various processes that are not explicitly listed in the following, as they do not have the objective of deriving tolerance values.In general, the research topic represents an important quality aspect for additively manufactured components. Thus, the current status of achievable geometrical accuracy must be examined and optimized in the future in order to make the serial production of components accessible to a larger number of industries.State of the artGeometrical deviations are unavoidable due to the physical manufacturing of parts and can generally be divided into four categories [5]. The four categories are dimensional deviations (two- point dimensions), form deviations (e.g.

The four categories are dimensional deviations (two- point dimensions), form deviations (e.g. roundness or cylindricity), positional deviations (e.g. perpendicularity or position) and surface deviations (e.g. surface roughness). For conventional manufacturing processes the achievable manufacturing accuracy is often defined in standards, e.g. general casting tolerances. The achievable tolerance values for AM are not standardized so far. Thus, the existing literature is used to give an overview about the manufacturing accuracy for Fused Deposition Modeling (FDM), Selective Laser Sintering (SLS) and Laser Beam Melting (LBM) in the following.The geometrical accuracy of the FDM process has already been investigated in various studies. IPPOLITO ET AL. found a deviation of up to +0.7 mm [6]. MAHESH ET AL. analyzed free- form surfaces and observed deviations from the nominal shape of up to +2.5 mm [7].

IPPOLITO ET AL. found a deviation of up to +0.7 mm [6]. MAHESH ET AL. analyzed free- form surfaces and observed deviations from the nominal shape of up to +2.5 mm [7]. The machine manufacturer STRATASYS advertises achievable tolerance values of ±0.127 mm or ±0.04 mm per mm [8]. Further investigations focus on the influence of machine parameters, such as shrinkage factors [9, 10]. The outcome shows shrinkage is the dominating factor for dimensional deviations in FDM. MOHAMED ET AL. showed a summary of the current activities to improve the geometric438accuracy in FDM [11]. MINETOLA ET AL. developed a benchmark part where different nominal lengths are considered according to DIN EN ISO 286-1. The results show that FDM can achieve ISO tolerance (IT) classes between 11 and 16. The investigation of the geometrical characteristics is carried out on cubes with an edge length of 10 mm. The cubes are made of explicitly specified ABS material, produced on a low-cost printer.

The investigation of the geometrical characteristics is carried out on cubes with an edge length of 10 mm. The cubes are made of explicitly specified ABS material, produced on a low-cost printer. A design of experiment (DoE) delivers optimized process parameters, e.g. smallest possible layer thickness for low dimensional deviations [12]. Another paper on dimensional accuracy in FDM was published by NANCHARAIAH ET AL. [13]. For the investigations, a test specimen with a constant dimension of 25.4 mm in length and 12.7 mm in diameter is used. The outcome of this research work demonstrates, that layer thickness and filament width affect the part accuracy significantly. HANSSEN ET AL. published a study on the achievable dimensional accuracy of the Stratasys Fortus 360mc and 400mc systems [14]. They advertise the considered systems with achievable tolerances of ±0.127 mm or ±0.0015 mm per mm depending on the nominal dimension.

They advertise the considered systems with achievable tolerances of ±0.127 mm or ±0.0015 mm per mm depending on the nominal dimension. The investigations were executed on three Fortus 400mc machines. The test specimen had a dimension of 127 mm x 76 mm x 14 mm and was manufactured using ABS-M30 [14]. Furthermore, LIENEKE ET AL. investigated dimensional tolerances for additive manufacturing for FDM and classified the process into the IT-classes 9 to 14 according to DIN EN ISO 286-1 [15].TANG ET AL. investigate the influence of process parameters on the accuracy of SLS. According to them, the geometric accuracy is mainly caused by the temperature distribution in the build chamber, the material shrinkage, the laser beam offset and the laser speed. After finding improved settings for the main influencing factors, the errors remain below ± 0.2 mm [16].

After finding improved settings for the main influencing factors, the errors remain below ± 0.2 mm [16]. WEGENER and WITT [17] also demonstrate that the influence of the temperature distribution within the building chamber affects the accuracy as well as the mechanical properties. They even proclaim that the temperature distribution is the main reason for a lack of reproducibility [16]. This statement can be clarified by the in-process temperature measurement by JOSUPEIT and SCHMID [18]. Further studies deal with shrinkage modelling to reduce the occurring deviations [19, 20]. RAGHUNATH, PANDAY and YANG ET AL. tested cuboids and show relations between material shrinkage and various process parameters [19, 20]. SENTHILKUMARAN ET AL. discussed the influence of building strategies on the accuracy of laser-sintered parts [21]. SEEPERSAD ET AL. showed manufacturing limitations for SLS.

SENTHILKUMARAN ET AL. discussed the influence of building strategies on the accuracy of laser-sintered parts [21]. SEEPERSAD ET AL. showed manufacturing limitations for SLS. The investigation deals with geometrical deviations on simple elements such as walls, holes, cylinders and complex elements such as gears [22]. However, SEEPERSAD ET AL. performed a qualitative assessment of the occurring deviation, but they did not mention numerical tolerance values. Furthermore, most of the literature references use only one nominal dimension for the executed investigations. LIENEKE ET. AL.

Furthermore, most of the literature references use only one nominal dimension for the executed investigations. LIENEKE ET. AL. [23] investigated the influence of the position, the alignment and the selected nominal dimension on the accuracy of the SLS on the geometric accuracy, whereby laser sintering was classified in IT classes 13 to 15 according to DIN EN ISO 286-1 under consideration of the boundary conditions [23, 24].COOKE and SOONS deal with deviations in the range of dimension, shape and position in the field of LBM and Electron Beam Melting (EBM). Based on a test specimen with a nominal dimension of 100 mm, dimensional deviations between -0.2 mm and +0.1 mm in the x-direction and -0.2 mm and +0.05 mm in the y-direction were defined. The circularity exhibits tolerance zones between 0.094 mm and 0.156 mm, which are calculated between the minimum and maximum deviation [25]. HANUMAIAH ET AL.

The circularity exhibits tolerance zones between 0.094 mm and 0.156 mm, which are calculated between the minimum and maximum deviation [25]. HANUMAIAH ET AL. investigated form and location deviations and derived tolerances for Direct Metal Laser Sintering (DMLS). A straightness tolerance of 0.0372 mm is defined in this publication. The deviations for flatness are examined within a tolerance of 0.0868 mm and the circularity tolerance is estimated to be 1.5320 mm [26]. Furthermore, LIENEKE ET. AL. [27] experimentally investigated dimensional deviations in LBM for different orientations and439nominal dimensions up to 80 mm and derived IT-classes between 12 and 17 according to DIN EN ISO 286-1.Concluding, the abovementioned literature demonstrates a large variation of observed geometrical deviations on different machines, test specimens and boundary conditions, which are mostly not listed in detail.

These differences can be explained by the large number of factors influencing the geometrical accuracy. Therefore, it is difficult to compare individual investigations with different processes, materials, machine settings and test specimens, if the manufacturing boundary conditions are not mentioned. This fact highlights that there is no generally known, reliable and comprehensive information about tolerances for AM processes. Thus, a uniform method needs to be developed to examine geometrical deviations and to derive realistic tolerance values. This publication is based on further results and extends an existing method [15, 23, 24, 27, 28].Purpose of investigationIn general, proven knowledge about the achievable manufacturing accuracy should be known for each manufacturing process in order to enable a solid component design and tolerancing. Furthermore, this allows to plan and to realize the development of the respective component process chain including further process steps.

Furthermore, this allows to plan and to realize the development of the respective component process chain including further process steps. This fact applies both to conventional manufacturing and to additive manufacturing processes. As already mentioned, the accuracy and achievable tolerance values are not yet comprehensively and publicly known for AM. For conventional processes, general tolerances are defined to the greatest possible extent, which express the manufacturing accuracy customary in the workshop and document it by means of a central entry on the technical drawing, e.g. DIN ISO 2768-1 and -2 for machining.Furthermore, no cross-process method is known that allows the investigation of the manufacturing accuracy. Therefore, the main objective of this investigation is the development of a method for the investigation of geometric deviations using standard workshop process settings, geometrically defined test specimens and selected measurement strategies.

The method is developed on the basis of the FDM, SLS and LBM processes. However, the method should be transferable to other additive or conventional manufacturing processes. The test specimens should cover different geometric shapes of measuring surfaces that allow the determination of geometric deviations. The classification of the geometries and their combinations is based on the definition of standard elements according to ADAM [29]. In the experimental investigations, the determination of geometric deviations and the quantification of the manufacturing accuracy by realistic tolerance values is focused. The results are described and interpreted within and across the considered processes.Methodical approachIn the beginning of the method development, necessary requirements are defined.

The results are described and interpreted within and across the considered processes.Methodical approachIn the beginning of the method development, necessary requirements are defined. Thefollowing requirements should be considered:Transferability: The methodical investigation of geometrical deviations should be applicable to different additive manufacturing processes as well as to conventional processes.Universality: The method should include technically relevant geometric elements and differently shaped measuring surfaces.Scalability: During the development of the method, the focus is on the FDM, SLS and LBM processes.

Nevertheless, the method should be adaptable to other manufacturing processes.440Thus, considered factors, for example the nominal dimension of test specimens, should be scalable.Transparency: The method is developed using standard workshop settings along the process and thus offers the possibility to compare process optimizations with the initial state.Measurability and repeatability: Suitable measuring strategies should be investigated for the defined test specimens to allow a repetitive execution.Quantification: The method and the included experimental investigation should enable the derivation of quantitative tolerance values.Classification and comparability: The determined tolerance values should be classified into standards and compared with conventional processes.After the definition of the main requirements, the general structure and the step-by-step procedure of the method is shown. In the following, the seven main steps are described with the upcoming results.

In the following, the seven main steps are described with the upcoming results. Step 1: Definition of target values: The aim of the method is the investigation of geometrical deviations. So, the four geometrical deviations (dimensional deviations, form deviations, positional deviations and surface deviations) are defined as the target values. In addition, profile deviations are considered. Step 2: Identification of factors influencing the manufacturing accuracy: A cause-effect diagram is developed for each of the processes. In Figure 1 the diagram for the LBM process is shown. In this paper, mainly factors in the area of “human" and "measurement" are considered in detail. Influences regarding the "machine", "method" as well as "material" are determined as constant in this publication, because these influences had already been considered and published by e.g.

Influences regarding the "machine", "method" as well as "material" are determined as constant in this publication, because these influences had already been considered and published by e.g. KNOOP and JOSUPEIT for FDM and SLS [30, 31].Figure 1: Cause-effect-diagram - influencing factors on the geometrical accuracy (LBM)441Step 3: Selection of relevant influencing factors Within this paper, the following influences are considered in detail:“human”: geometry, dimension, part position in the build chamber, and orientation “measurement”: measuring equipment and measuring method (extraction strategy)Step 4: Definition of the variation range for selected influencing factors For each selected factor, a variation range is determined in order to investigate the influence of the factor.

According to the selected factors in this publication the following ranges of variation for the experimental investigation are defined:“human”: o The geometry is subdivided into three main groups according to the definition of standard elements by ADAM [29]. For the experimental investigation the measuring areas of the test specimen are defined as planar, cylindrical, and spherical surfaces.Figure 2: Classification of the geometry – planar, cylindrical and spherical areaso The dimension (e.g. length, width, height or diameter) of the different geometries and measuring areas are set according to the DIN EN ISO 286-1. The variation range of the nominal dimensions is documented in Table 1.

length, width, height or diameter) of the different geometries and measuring areas are set according to the DIN EN ISO 286-1. The variation range of the nominal dimensions is documented in Table 1. Although the standard offers higher nominal dimensions, the value of 500 mm is defined as the maximum by the build chamber limits of the considered processes.Table 1: Selected nominal dimension according to DIN EN ISO 286-1Nominal dimension / mm3 6 10 18 30 50 80 120 180 250 315 400 500 600 …o The chosen position in the build chamber is an often-mentioned influencing factor in literature [29, 30, 31]. In order to analyze the possible influence on the geometrical accuracy in detail, different numbers of positions are investigated in the x-y plane depending on the nominal dimension and build chamber space.

Thus, in dependence of the selected process between 81 and 225 positions are investigated for small nominal dimensions (3 to 18 mm), nine different positions for nominal dimensions up to 80 mm and three positions for nominal dimensions between 120 mm and 500 mm. Figure 3 demonstrates the nine different positions in the x-y plane that are considered up to 80 mm for all processes.442Figure 3: Schematic representation of the nine positions in the x-y plane with cuboidso Another important influencing factor is the orientation of test specimen. Due to this, the factor is investigated by the following variation range depending on the target value. Dimensional deviations are recorded in x, y and z alignment in order to have the possibility to optimize the shrinkage factors along the axis. Form and positional deviations are investigated in 0°, 30°, 45°, 60° and 90° orientation.

Form and positional deviations are investigated in 0°, 30°, 45°, 60° and 90° orientation. Because literature demonstrates that the scattering in the surface roughness is very high, for example in SLS [31], the measuring areas for surface deviations are orientated in 0°, 30°, 45°, 60°, 90°, 120°, 150° and 180°. For further influencing factors, the variation ranges were defined, but they are not pictured in this paper. For instance, for the process-specific influencing factor “layer thickness” that appears in the category “method – machine parameters” (s. Figure 1) is varied between 0.06 mm, 0.120 mm and 0.180 mm for SLS, between 0.030 mm and 0.050 mm for LBM and between 0.127 mm, 0.178 mm and 0.254 mm for FDM.Step 5: Design and classification of test specimensBased on “Step 4”, test specimens that allow the consideration and the experimental investigation of the defined influences within the belonging ranges are designed.

Due to the definition of the geometry, test specimens with planar, cylindrical and spherical measuring areas are developed and selected. Table 2 shows the classification of test specimens as well as the investigated target values on each test specimen design.

Table 2 shows the classification of test specimens as well as the investigated target values on each test specimen design. Each test specimen design exists in an “outer” and an “inner” design, in order to investigate outer and inner diameters of cylinders and holes, because previous studies have shown that inner and outer dimensional deviations can vary significantly for the same nominal dimension.443Table 2: Classification of test specimen designs according to the geometry of measuring area and the targets values: dimensional (two-point measurement), form (flatness, straightness, roundness, cylindricity), profile (line and surface), positional (direction, location and run) and surface deviationsAs an example of the above-mentioned classification and the demonstration of the variation range of influencing factors, Figure 4 demonstrates the application of the orientation for cuboids and hollow cylinders with fixed dimensions.

In addition, the dimensions are varied according to the nominal dimension of DIN EN ISO 286-1 (s. Table 1).Figure 4: Representation of cuboids and hollow cylinders in five different orientations444Step 6: Definition of tactile measuring methodsAfter the definition of the different test specimen designs, the measuring equipment and measuring strategies are selected. The following topics belong to the point “measurement”, “measuring equipment” and “measuring method (extraction strategy)” within the cause-effect- diagram in Figure 1.The Nikon Altera 8.7.6 coordinate measuring machine is used to record the target geometrical deviations. The tactile measuring system has a volumetric accuracy of 1.8 µm + N/400 (N: nominal dimension). For the tactile measurement a probe with 40 mm length and ruby sphere diameter of 2 mm is used.The planning of the measurement method for each test specimen design is important to achieve realistic and meaningful results.

For quality standardization and the industrial verification of conformity, DIN EN ISO 17450-1 and -2 are adopted. The standard deals with the verification of the specification of parts to be evaluated. The connection between the extraction and association of the part is covered by DIN ISO 14660. The extraction is a derivation of the real part geometry into the geometry of a grid of measuring points. The determination of the position and number of measuring points is relevant for the degree of approximation of the real geometry [33]. Thus, for each test specimen, a meaningful extraction and association strategy is selected and developed. Especially in AM, the inhomogeneous surface roughness on a single component complicates the measurement of realistic values. Because of this, different extraction strategies are experimentally analyzed and evaluated for planar, cylindrical and spherical geometries.

Because of this, different extraction strategies are experimentally analyzed and evaluated for planar, cylindrical and spherical geometries. Within this publication the experimental preliminary investigation for planar and cylindrical geometries are considered in detail.Planar geometries – cuboids: Planar surfaces are tested with three different extraction strategies. Starting from a nominal area of 10 mm by 10 mm to be examined, triangular grids, rectangular grids and the random distribution of points (by CMM software) are used. As the nominal length increases, so the number of measuring points increases proportionally. Since in preliminary studies the recommended numbers of measurement points (e.g. nine measurement points for a plane) showed non-realistic deviations, the minimum numbers of measurement points on Level 1 were defined to a value approximately twice as large. The three extraction strategies are varied in three levels of measuring points (s.

The three extraction strategies are varied in three levels of measuring points (s. Figure 5): o Level 1: Low number of measuring points o Level 2: Medium number of measuring points o Level 3: High number of measuring points Figure 6 shows the results for the flatness deviation on the SLS cuboid with a planar measuring area of 10 mm by 50 mm that was manufactured in 0° orientation. Each measuring strategy was performed three times on the same test specimen. In sum, each test specimen was manufactured three times. Exemplary, the diagram shows the minimum, maximum and the average value of three measurements on one test specimen.

In sum, each test specimen was manufactured three times. Exemplary, the diagram shows the minimum, maximum and the average value of three measurements on one test specimen. The 0° orientation was used, because this orientation causes the highest flatness deviations due to the higher exposure area in x-y plane and the resulting warpage of the test specimen.445Figure 5: Extraction strategies for planar measuring areas with three different levels of pointsIt can be seen that the extraction strategy “random distribution of points (RD)” determines significantly smaller deviations than the other two strategies for the examined surface, regardless of the number of measuring points. With an increase in the number of measuring points, the recorded flatness deviations of the RD strategy increase. The triangular grid (TG) and rectangular grid (RG) strategies record almost equivalent flatness deviations at Level 2 (TG:38 and RG:45) and Level 3 (TG:88 and RG:80).

The triangular grid (TG) and rectangular grid (RG) strategies record almost equivalent flatness deviations at Level 2 (TG:38 and RG:45) and Level 3 (TG:88 and RG:80). In these ranges, the strategies are not significantly different from each other. However, in the case of the rectangular grid with a low number of measuring points (RG:20), much smaller deviations in flatness are recorded compared to the triangular grid (TG:23).

However, in the case of the rectangular grid with a low number of measuring points (RG:20), much smaller deviations in flatness are recorded compared to the triangular grid (TG:23). The average value of the deviations for the rectangular grid is 0.068 mm and for the triangular grid it is 0.121 mm on level 1.Influence of the extraction strategy on the flatness deviation – SLS cuboid L50x10 in 0° orientation0.150m m/ n o i t a i v e d s s e n t a l f0.1000.0500.000TG:23 RG:20 RD:23TG:38 RG:45 RD:45TG:88 RG:80 RD:88number of measuring points / -minimum minimum minimummaximum maximum maximumTriangular grid Rectangular grid Random distributionFigure 6: Influence of the extraction strategy on the measured flatness deviation446Cylindrical geometries – hollow cylinders: Cylindrical surfaces are measured with two different extraction strategies. Firstly, roundness profiles (RP) and secondly, shifted roundness profiles (SRP) are examined.

Firstly, roundness profiles (RP) and secondly, shifted roundness profiles (SRP) are examined. The two strategies are identical in the number of measurement points as well as the measuring cycles for each variation level (s. Figure 7). Only the distribution of the measurement points is changed here. In general, the number of measuring points is adjusted via the circle diameter of the cylindrical geometries. So, the both strategies and their four levels of measuring point numbers are demonstrated on the hollow cylinder with a nominal diameter of 30 mm and a nominal length of 30 mm (Figure 7).Figure 7: Extraction strategies for cylindrical measuring areas with four different levels of pointsFigure 8 shows the deviation of the cylinder form of the SLS hollow cylinders with a length of 30 mm and a diameter of 30 mm (L30xD30) in 0° orientation. The average value is based on three measurements on one test specimen (Figure 8). Each test specimen was manufactured three times.

The average value is based on three measurements on one test specimen (Figure 8). Each test specimen was manufactured three times. The recorded average values of the cylindricity deviations is shown as a function of the extraction strategy and the number of measuring points. The red curve stands for the roundness profile strategy (RP) and green curve for the shifted roundness profile strategy (SRP). Comparing both curves, it can be clearly seen that different averaged measured values of the cylinder form deviation are achieved with a low number of measuring points.For example, with a minimum number of measuring points (8 points), the measured value of the roundness profile is 0.048 mm and the measured value of the shifted roundness profile is 0.102 mm. As the number of measuring points increases, the recorded deviations in cylindricity also increase. The more points are used to measure the cylindricity, the closer the measured mean values are to each other.

The more points are used to measure the cylindricity, the closer the measured mean values are to each other. With 675 measuring points, both extraction strategies have approached a measured value of 0.268 to 0.275 mm. For all measurements it is obvious that the scattering of the three repeated measurements is very small (Figure 6 and 8).447Figure 8: Influence of the extraction strategy on the measured cylinder form deviationBased on the experimental preliminary investigation, hints for the measuring method are derived. The recommended extraction strategies strike a balance between accuracy and reliability, measurement time and economy of measurement results:1. Note: Full-surface measurement point distribution and choice of a structured distribution strategy is needed.2. Note: Sufficiently high number of measuring points:a. For planar measuring areas: Triangular grid for detection is useful. At least 3 measuring lines and 5 measuring points on a 10 mm by 10 mm surface.

For planar measuring areas: Triangular grid for detection is useful. At least 3 measuring lines and 5 measuring points on a 10 mm by 10 mm surface. The number of measuring points increases proportionally to the nominal size of this surface (s. Figure 5).b. For cylindrical areas: Shifted roundness profiles or similar detection strategies are recommendable. Cylinders with a nominal length of 10 mm and a diameter of 6 mm should be measured with at least 3 measuring cycles and 10 measuring points per cycle. Here, the number of cycles and measuring points per cycle increases with increasing nominal dimension (s. Figure 7).3. Note: Selection of a suitable probe diameter: Recommendation of small probe diameter for planar, cylindrical and spherical measuring areas.448Step 7: Definition of constant boundary conditionsIn the last step of the method development, the test plan for the experimental investigation and the boundary conditions are set up.

All manufactured test specimens are conditioned in standard climate according to acceptance regulations of plastic and metal components (DIN EN ISO 291). Furthermore, Table 3 lists the boundary conditions for the AM processes that are used for the manufacturing of the test specimens.

Furthermore, Table 3 lists the boundary conditions for the AM processes that are used for the manufacturing of the test specimens. These settings are defined as constant and are used as a standard during the method development and validation.Table 3: Boundary conditions for FDM, SLS and LBMprocess factorFDMSLSLBMmachineStratasys Fortus 400mcEOS EOSINT P396SLM Solutions SLM 280 HL 1.0Layer thickness178 µm120 µm50 µmshrink factors (x/y/z)0.55 % 0.55 % 0.59 %3.2 % 3.2 % 2.55 - 1.4 %0.223 % 0.223 % 0.223 %materialABS M30PA2200 (50%-50%)316LsupportSR-30Dispers powderSolid block supportpost-processingMechanical/chemical removal of the supportGlass ball blastingExperimental investigationWithin this publication, the results for two selected test specimens and selected influencing factors are presented. For this purpose, the above-mentioned cuboids and hollow cylinders are focused.

For this purpose, the above-mentioned cuboids and hollow cylinders are focused. On the basis of the cuboids, the possible influence of the position on the dimensional deviations in LBM is considered. Therefore, cuboids with a constant nominal dimension of 12 mm x 12 mm x 12 mm are used, which are manufactured on 81 positions in the build chamber. The dimensional deviations of the test specimens are measured in the x-y plane. The hollow cylinders are manufactured with different diameters and lengths and in the orientations 0°, 45° and 90°. In addition to the dimensional deviation, the cylindrical form deviation of the hollow cylinders is determined according to the defined measuring method. In the following, the two test specimens and the results are presented in more detail. Cuboids – planar measuring areas:The following results are shown for LBM. Table 4 sums up the considered influencing factors and the variation range of each factor that is analyzed.

Cuboids – planar measuring areas:The following results are shown for LBM. Table 4 sums up the considered influencing factors and the variation range of each factor that is analyzed. The investigation focuses on the influence of the chosen position in the LBM building chamber. The dimensional deviations of the manufactured cuboids are measured in the x-y plane. The entire build job (s.

The dimensional deviations of the manufactured cuboids are measured in the x-y plane. The entire build job (s. Table 3) is repeated three times.449Table 4: Design of experiment for the investigation of the influence of the position in the LBM build chamberdrawingtarget valueinfluencing factorvariation rangedimensional deviationtest specimencuboidsgeometryplanarnominal dimension12 mmposition in build chamber81 positionsorientation90°measuring method: coordinate measuring machine – triangular grid (TG) – level 3schematic representation: build job design and the manufactured test specimens on the build platformAfter the measurement of all cuboids, the dimensional deviation is calculated for each position. The results are shown in Figure 9. The diagram presents the mean values of three build jobs. The occurring dimensions emphasize that the measured dimensions increase in the outer areas of the build platform, especially in the corners.

The diagram presents the mean values of three build jobs. The occurring dimensions emphasize that the measured dimensions increase in the outer areas of the build platform, especially in the corners. It is obvious, that the extreme positions in the four corners show the highest actual dimensions and so the highest dimensional deviations. The actual dimensions vary in the corners between 12,11 and 12,16 mm. The middle position shows an actual dimension of 12,05 mm. In contrast to further investigations [29], the chosen position shows an influence on the geometrical accuracy. Similar results had already been shown for the SLS process [23, 24, 30]. However, in SLS the deviations occur due to the inhomogeneous thermal situation in the build chamber during and after the building process [30]. Such a high influence due to the temperature in LBM is not expected.

Such a high influence due to the temperature in LBM is not expected. Because of this, some feasible explanations are pointed out hereafter.450Figure 9: Dependence of the occurring dimensions as a result of the selected position in LBMThe cause of these position-specific deviations cannot be fully explained. It seems unlikely that thermal effects such as in the SLS process, where the different temperature distribution in the build chamber causes the plastic components to shrink inhomogeneously, are the main influence. A possible explanation for the increased deviations in the corners of the LBM chamber could be the laser beam guidance. By deflecting the laser beam into the corners, an elliptical change in the laser spot occurs without the usage of any correction systems. However, this change is prevented by F-theta lenses or dynamic systems as standard in LBM machines. So, one reason could be an improperly calibrated component in the beam guidance.

However, this change is prevented by F-theta lenses or dynamic systems as standard in LBM machines. So, one reason could be an improperly calibrated component in the beam guidance. Furthermore, it is conceivable that there is an increased influence on the gas flow in the corners, for example due to the position of the recoater. An insufficient gas flow velocity could not be removed possible welding spatter that could lead to increased dimensional deviations. Last, the left-hand side of the build chamber (s. Figure 9) also shows slightly increased actual dimensions in contrast to the right-hand side, which could be possible due to an increased weld spatter density in this area.Hollow cylinder – cylindrical measuring areas:The investigation of hollow cylinders is shown in the following for LBM. Table 5 summarizes the influencing factors and the variation range of each investigated factor. The investigation considers on the geometrical deviations on cylindrical measuring areas.

Table 5 summarizes the influencing factors and the variation range of each investigated factor. The investigation considers on the geometrical deviations on cylindrical measuring areas. The focus here is on the dimensional and form deviations on the outer diameter of the test specimen. Further considered deviations are not presented in this paper. The hollow cylinders are built in each influencing factor variation three times. Due to solid support material, the measuring method need to be adapted locally [28].

Due to solid support material, the measuring method need to be adapted locally [28]. Measuring points were omitted at surfaces with solid support material appearing in 0° and 45° orientation.451Table 5 Design of experiment for the investigation of geometrical deviations in LBM using hollow cylindersdrawing / target valuesinfluencing factorvariation rangedimensional, form and positional deviationstest specimenhollow cylindergeometrycylindricalnominal outer diameter d10, 18, 30, 50 mmnominal length l10, 18, 30, 50 mmposition in build chamberfreely chosenorientation0°, 45°, 90°measuring method: coordinate measuring machine – shifted roundness profiles (SRP) – level 4schematic representation: build job design and real test specimens in 45° and 90° orientationThe hollow cylinders are measured with the coordinate measuring machine using the above- mentioned measuring method.

In 0° and 45° orientation some surfaces need solid support material in order to guarantee a robust manufacturing process (s. Table 5). Those areas are left out during the measurement. This emphasizes that especially for LBM the defined measuring method needs to be adapted locally. However, it is obvious that surfaces with solid support material need to be machined for the support removal. The results for the occurring dimensional deviations are shown in Figure 10. The diagram presents the mean, maximum and minimum values of the dimensional deviations. Both the chosen orientation and the dimension in diameter and length show a significant influence on the dimensional accuracy. In general, the occurring deviation increases with increasing nominal length and diameter independent of the selected orientation. For higher dimensions, for example L50xD50, the influence of the chosen orientation is strengthened.

For higher dimensions, for example L50xD50, the influence of the chosen orientation is strengthened. On average, the 90° orientation shows the smallest dimensional deviations and also the scattering of the measured values is the smallest. In contrast, the 0° orientation mostly indicates the largest deviations. In particular, larger nominal diameters show a larger undersize and the highest scattering values in this orientation.4520.400.300.200° Average 45° Average 90° Average0° Min. 45° Min. 90° Min.0° Max. 45° Max. 90° Max.m m0.10/ n o i t a i v e d r e t e m a i d0.000.100.200.300.400.500.60description and nominal dimension (length x diameter) / -Figure 10: Dimensional deviations of the outer diameter in dependence of the dimension and orientationThe explanation on dimensional deviations can also be transferred to the cylindricity of the outer surface of the cylinders.

The results in Figure 11 show that the cylinder form deviations on 90° orientated cylinders is the smallest in average. Similar to the dimensional deviations, cylinders in 0° and 45° orientation show the highest cylinder shape deviations. This distinction becomes clearer when the diameter and the length of the cylinder increase. Furthermore, the scattering of the measured values predominantly raises with a decreasing orientation.m m/ n o i t a i v e d m r o f1.000.900.800.700.600.500.400.300° Average 45° Average 90° Average0° Min. 45° Min. 90° Min.0° Max. 45° Max.

45° Min. 90° Min.0° Max. 45° Max. 90° Max.r e d n0.20i l y c0.100.00description and nominal dimension (length x diameter) / -Figure 11: Cylinder form deviation of the outer diameter in dependence of the dimension and orientation453The resulting deviation can be explained on the one hand by the scaling of the components with standard shrinkage factors independent on the actual dimension of the test specimen and on the other hand by the induced residual stresses, which are induced significantly higher in the cylindrical measuring surface in the orientations 0° and 45°. The orientation 90° shows the lowest deviations because the respective layers are manufactured exactly on top of each other, which means that no free overhangs have to be built into the powder bed.

The orientation 90° shows the lowest deviations because the respective layers are manufactured exactly on top of each other, which means that no free overhangs have to be built into the powder bed. In addition, test specimens in 90° orientation are not influenced by the stair-stepping effect.Summarizing it can be said that the selected influencing factors and their variation ranges are fitting so far. Furthermore, it was necessary to carry out preliminary investigations in terms of suitable measurement methods in order to ensure a realistic representation of the occurring deviations. Relative to the geometric influencing factors, significant dependencies on the position in the build chamber, the orientation and on the dimension of the test specimens could be demonstrated.

Relative to the geometric influencing factors, significant dependencies on the position in the build chamber, the orientation and on the dimension of the test specimens could be demonstrated. In addition to the test specimens and deviations shown here, all test specimens shown in Table 2 were investigated experimentally for the FDM, SLS and LBM processes using adapted measurement methods.Conclusion and outlookAdditive manufacturing provides new technical and economical capabilities for the product development by contrast to established manufacturing processes. However, many process specific challenges hamper the usage of AM for serial production. One of the biggest challenges is the insufficient geometrical accuracy. Since the achievable manufacturing accuracy of the processes is often unknown, a successful post-processing is also difficult to consider before manufacturing.

Since the achievable manufacturing accuracy of the processes is often unknown, a successful post-processing is also difficult to consider before manufacturing. The present publication provides a contribution to the systematic investigation of the geometrical accuracy in AM. On the basis of relevant influencing factors and variation ranges, test specimens and suitable measuring methods were systematically developed and evaluated. The resulting method was tested by experimental investigations. The defined influencing factors and their variations could be considered as suitable. The defined measurement method, which describes in particular the extraction strategy, was also considered as sufficient. Besides tactile measuring methods with a high number of points, optical measuring methods offer a clear advantage in order to investigate the local geometrical deviations and their causes in detail.

Besides tactile measuring methods with a high number of points, optical measuring methods offer a clear advantage in order to investigate the local geometrical deviations and their causes in detail. Especially due to local warpage and varying surface roughness on a single test specimen, a sufficient number of measuring points has to be provided. In addition, the number of points and the strategy are dependent on the geometry of the measuring area. The results illustrate an influence of the position, orientation and the selected nominal dimensions on the dimensional and form deviations of additively manufactured parts quantitative. The reasons for the occurring deviations could not yet be fully clarified within this work. In addition to geometric influences, process-related influences and machine components must also be taken into account in more detail in the upcoming steps. These influences enable to improve the geometric accuracy of the different processes.

OMETRY AND PROCEDURE FOR BENCHMARKING SFF AND HYBRID FABRICATION PROCESS RESOLUTIONVito R. Gervasi, Adam Schneider, Joshua Rocholl the Milwaukee School of Engineering, Milwaukee, WisconsinABSTRACTSince the advent of SFF and RP a number of SFF benchmarking geometries and methodologieshave been developed and employed with some similarities but limited standardization. Minimal information has been published in regard to a standard method of measuring the resolution limits or capabilities of SFF and SFF-based hybrid processes. In an effort to benchmark resolution limits of SFF and Hybrid Fabrication processes, several benchmarking geometries were developed to capture the resolution capabilities, specifically hole size and rod size range, of multiple hybrid fabrication path steps and a hybrid path as a whole. These useful geometries are shared with the SFF community and procedures for their use are described in this paper.1.

These useful geometries are shared with the SFF community and procedures for their use are described in this paper.1. INTRODUCTIONA combination of procedure and benchmark geometries was developed for evaluating hybridfabrication paths. This benchmarking approach is intended for, but not limited to, design parameter input of manufacturable optimized intertwined lattice structures, made up of rod and hole elements. When designing objects using an optimized intertwined lattice structure approach, sub-surface structure consists of two or more distinct phases. These phases are in the form of an intertwined, interconnected, lattice structure with a range of rod and hole sizes as well as a range of aspect ratios.

These phases are in the form of an intertwined, interconnected, lattice structure with a range of rod and hole sizes as well as a range of aspect ratios. Before designing subsurface lattice structures much must be known about the manufacturability of these rod and hole elements if they are to be successfully fabricated using SFF or SFF-based hybrid fabrication techniques.PurposeFor optimized intertwined lattice structures to be designed for manufacturability, hybrid pathcapabilities and limitations must be well understood. In fact, before designing an optimized lattice structure, which may approach one million or more rod and hole elements, manufacturing process capabilities are required as a design input. In other words, the range of rod and hole sizes feasible, with a given hybrid fabrication path, must be known and provided on the input side of the design process.

In other words, the range of rod and hole sizes feasible, with a given hybrid fabrication path, must be known and provided on the input side of the design process. Other applications employing lattice structures or ultra fine features are in need of SFF process and hybrid path evaluation before widespread use and measurable quality improvements can take place. This paper shares a combination of procedure and benchmark geometries used for evaluating hybrid fabrication paths at MSOE. This combination is intended for, but not limited to, the design of manufacturable optimized intertwined lattice structures, made up of rod and hole elements.ScopeThe objective of this benchmarking approach is to determine the rod and hole size-range capabilities and aspect ratio limitations of individual steps of a hybrid fabrication path and the path as a whole. The resulting data provide critical input parameters for creating manufacturable designs.

The resulting data provide critical input parameters for creating manufacturable designs. Benchmarking configurations with integrated rod and hole diameters were considered as well as daisywheel configurations but the effort of evaluating these complex objects and the potential for unnoticed defects steered geometries to simpler forms.To develop the benchmarking geometries shown herein three criteria were applied with simplicitybeing a guiding factor:(cid:1) (cid:1) (cid:1)covers one order of magnitude in feature sizes relatively easy to use and evaluate captures aspect ratios if needed493Beyond the scope of this initial effort, though important, were material properties and internaldefect considerations.Figure 1.

These phases are in the form of an intertwined, interconnected lattice structure made up of rods and holes with a range of feature sizes and aspect ratios (figure 1).

To design these subsurface lattice structures much must be known about the fabrication method and manufacturability of these rod and hole features.A literature search provided the following list of process benchmarking, calibration, or evaluationmethods, none of which specifically address rod and hole resolution capability:(cid:129) 3D Systems WindowpanesTM [3] (cid:129) 3D Systems ChristmasTreesTM [3]494(cid:129) A number of RP benchmarking parts have been proposed and studies have been presented whichfocus on amplifying warpage [4](cid:129) Detecting cure-through [5] (cid:129) Determining thermal gradient impact on dimensions [6] (cid:129) Providing overall comparisons of RP process capabilities (cid:129) Chrysler [7] published a study comparing the top RP processes in 1993, using a speedometer(cid:129)adapter-comparison part Jacobs suggests the presence of noise in all processes involving a phase change, calling it a (cid:1)random noise shrinkage constant.(cid:1) [8]Overall, RP-related publications were not directly helpful but did provide much usefulinformation on sources of error and benchmark design approaches.2.

PROCEDUREThe procedure for evaluating a single step of a hybrid path is illustrated in the flow chart shown in figure 2. The flowchart is applied to determine the rod and hole resolution capability of each process step as well as aspect ratio limitations. Up to four of six benchmarking geometries are employed depending on the type of process and hole or rod form. Upon completion of the flowchart for each process step, the step specific data are stacked to determine an overall hybrid path capability. If only SFF is used to generate the object for a given application, only one pass through the flowchart is required. Aspect Ratio Sensitivity?For some processes, aspect ratio sensitivity is very critical. For example, if a ceramic rod isformed with a high length-to-diameter ratio, the ceramic feature may be easily damaged duringStart StartFigure 2. Flowchart and Figure 2.

SFF and SFF-based hybrid fabrication path resolution flowchart.495processing. If an expendable pattern is used to initially form the feature the pattern material may expand due to moisture (from the ceramic mold slurry) or the pattern may thermally expand (during pattern burnout) damaging the ceramic mold feature in either case. In other cases, such as the selective laser sintering process or metal casting of rod features, aspect ratio is not as problematic since both processes are capable of large aspect ratios. In the case of stereolithography, the maximum unsupported feature length is well known by most machine operators, driving the maximum rod length. If the process step of interest is sensitive to aspect ratio, the BMP1AR is recommended. There are two forms of the BMP1AR and the user must determine which form will result in the desired feature shape, rod or hole, through the required number of transfers.

There are two forms of the BMP1AR and the user must determine which form will result in the desired feature shape, rod or hole, through the required number of transfers. It may take more than one transfer to test a particular step so the user may start with the same form that they end with. The BMP1-AR(cid:1)s are intended to uncover any aspect ratio limitations of a particular process step. The form of BMP1-AR is shown with 1:1, 2:1, 3:1, 5:1, & 10:1 ratios and can be scaled in the Z to provide higher aspect ratios. The BMP rung diameters range is size one order of magnitude and can be scaled. When generating BMP(cid:1)s to evaluate post SFF steps it is recommended that expendable patterns be produced with features 2-3 rung smaller that the process step is thought to be capable of. If it is not the SFF process being evaluated the pattern can be generated using the best build angle possible.

If it is not the SFF process being evaluated the pattern can be generated using the best build angle possible. Regarding the basic BMP1 design, each test feature diameter is 10 percent smaller than the largest feature in a linear fashion down to the smallest feature who(cid:1)s dimensions are 10 percent of the largest feature. Therefore, if the larger feature is scaled to 1 mm, the feature sizes would proceed from largest to smallest as follows: 1.0mm, 0.9mm, 0.8mm, 0.7mm, 0.6mm, 0.5mm, 0.4mm, 0.3mm, 0.2mm, to 0.1mm for the smallest rung. By using a range of sizes the (cid:1)step-cutoff(cid:1) or step limit can be brought into focus relatively quickly. The part could be modified to focus on a smaller range if needed.If the aspect ratio for the rods or holes is not an issue the user proceeds to one or both BMP1(cid:1)s.BMP1 is a simple single aspect ratio latter-like form with 10 rungs, each consecutively 10 percent smaller than the largest.

The two forms of the part are holes and rods. The use is very straightforward. Set the smallest feature 2-3 rungs smaller than the expected smallest feature and process the object through the needed step. Build BMPs at the (cid:1)worst-case(cid:1) build angle. There are two forms of the BMP1 and the user must determine which form will result in the desired feature shape, rod or hole, for the desired step, through the required number of transfers. It may take more than one transfer to test a particular step so the user may start with the same form that they end with. It is recommended to start with both forms and follow through all steps to get a complete picture of process capability. 5 Copies are recommended for each step.Additive Process?If the step being evaluated is an additive process it is important to verify that all build angles are producable. BMP2+/- are used to verify that the user does not overlook a problematic build angle and to verify the results of BMP1.

BMP2+/- are used to verify that the user does not overlook a problematic build angle and to verify the results of BMP1. BMP2+/- are scaled to match the minimal hole and rod diameter revealed with BMP1+/-. If BMP2+/- is unable to build features at a particular angle it is recommended to repeat the flow chart using the new (cid:1)worst-case(cid:1) build angle. All features must form to pass the verification. Machine variation from build to build may play a role in BMP2 failure. At least five copies are recommended. If aspect ratio results are a concern from BMP1 aspect ratios on BMP2 can be adjusted.Additional notes and suggestions:(cid:129) Patterns used to evaluate subsequent steps can be built at (cid:1)best-case(cid:1) build angle to provide databeyond the (cid:1)current(cid:1) capability of SFF resolution.(cid:129) Steps may be combined if they are inseparable. Ideally each step is evaluated independently of(cid:129)all other steps.

Ideally each step is evaluated independently of(cid:129)all other steps. If SFF support structure is required on a rod or hole element it is not a feature that can be produced as a lattice element.4964973. EXAMPLE RESULTS (cid:129) A graphic is prepared for each process step as shown in graph 2. Those features formed are shown in green. Features that did not form or defective features are shown in yellow.

Combined Hybrid Path Results, Rod FeatureCAUTIONARY MEASURES AND CONSIDERATIONS(cid:129) Build Crash Misinformation -if small features break free during the SFF process they maydamage features during recoats that would otherwise be successfully formed.(cid:129) SFF Z Error -error due to material being added to down-facing SFF surfaces may be misleadingand CAD compensation may close off holes unintentionally.(cid:129) Laser Beam Compensation Error (cid:1)Some features may not be included in a slice file due to laser beam compensation while other features may be larger than CAD due to laser beam diameter error.(cid:129) Data-use risk- When using the resulting data for design input for any application the designer must understand the risk involved and the probability of features forming as expected.

For critical aerospace or medical applications sufficient safety factors, based on sound statistical data must be applied.(cid:129) Stair Stepping Error- Stair-stepping can lead to variations in effective diameter as rods becomesmaller. Stair-stepping may also present notch sensitive regions not to be overlooked. (cid:129) Other Properties- The aforementioned approach only addresses geometric capabilities of aprocess. Additional testing is required to characterize other critical properties.5. CONCLUSIONSAn approach and benchmark geometries were developed for use in evaluating the (cid:1)rod(cid:1) and(cid:1)hole(cid:1) capability of SFF and SFF-based hybrid fabrication paths.This new approach is being used for, but is not limited to, the design of optimized intertwinedlattice structures.This approach may be applicable to evaluating objects on a range of scales from sub-nano tomeso-scale.4984.6.

FUTURE RECOMMENDATIONS A concern not addressed with the six aforementioned benchmark geometries is the (cid:1)depth(cid:1) capability of a lattice field produced by a given SFF-based hybrid path. Benchmarking Geometry to capture lattice depth capability, driven by the results of BMP1 and BMP2, should be developed. fabrications processes could be developed based on BMP1 by scaling the form in one direction.7. ACKNOWLEDGMENTSThe authors acknowledge the financial support for the research by National Science Foundation(DMI 0140717).8. REFERENCES 1. Gervasi, R.R., Crockett, R.S., (cid:1)Composites with Gradient Properties from Solid Freeform Fabrication,(cid:1) Solid Freeform Fabrication Conference Proceedings, Austin, Texas, August 1998.2. Stahl, D. C. and Gervasi, V. 2003. (cid:1)Design and fabrication of components with optimized lattice microstructures,(cid:1) Proceedings: 2003 NSF Design, Service, and Manufacturing Conference, Birmingham, Alabama, January, 2003.3.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperGeometry Limitations in Indirect Selective Laser Sintering of Alumina D. Sassaman1*, M. Ide†, J. Beaman*, C. Seepersad*, and D. Kovar*Center for Additive Manufacturing and Design Innovation and Walker Department of Mechanical Engineering The University of Texas at Austin, Austin, TX, 78712†ExxonMobil Research and Development Company Annandale, NJ 08801AbstractCeramics containing open channels with complex geometries can be manufactured by additive manufacturing (AM) and are of great interest in clean energy technologies. However, design limitations and guidelines for manufacturing these architectures with AM have not yet been established. In this work, we compare previously proposed geometry limitations for polymer selective laser sintering (SLS) to the geometries produced using indirect SLS in alumina.

In this work, we compare previously proposed geometry limitations for polymer selective laser sintering (SLS) to the geometries produced using indirect SLS in alumina. We focus on a subset of model shapes that are simple to produce and measure. We show that these rules provide a starting point for the design and manufacture of ceramic geometries using indirect SLS. However, there are additional considerations for AM of ceramics by indirect SLS that further limit the geometries that can be produced.1. IntroductionIndirect selective laser sintering (SLS) of ceramics uses a sacrificial polymer binder thatis mixed with a ceramic powder before processing. Indirect SLS simply replaces the consolidation/forming step in traditional ceramic processing; the same pre- and post-processing steps are required (Figure 1). In the case of indirect SLS, this consolidation step is comprised of two sub-processes; powder layer deposition and laser sintering.

In the case of indirect SLS, this consolidation step is comprised of two sub-processes; powder layer deposition and laser sintering. The spreading mechanism deposits a thin layer of a ceramic/binder mixture. During laser irradiation, only the binder melts, forming bridges between the ceramic particles. In this paper, to avoid confusion, the term “sintering” is defined as the act of joining particles together during the SLS process, and “densification” is defined as joining the ceramic particles during post-processing.Figure 1: Top) Traditional ceramic processing, Bottom) Indirect SLS ceramic processing1 Corresponding author contact information: doug.sassaman@utexas.edu688Complex geometries produced by indirect SLS have been studied in the context of cellular ceramics, metamaterials, flow reactors, and tissue scaffold engineering [1], but rules and limitations for designing them have not been developed.

Existing literature primarily focuses on geometrical accuracy of relatively simple shapes [2]–[6]. The most recent and comprehensive work on this topic comes from researchers at KU Leuven [3] and The University of Missouri Rolla [7]. For example, Nolte et al. were able to produce holes with relatively complex paths and straight holes with diameters of 1 mm ± 0.12 mm [7]. Nissen et al. proposed a methodology for determining geometry limitations in helical glass channels produced by indirect SLS [8]. Allison et al. developed geometry limitations, and a method for determining them, for polymer SLS [9]– [11]. Additionally, Allison et al. developed a metrology part, from which the geometries in this work have been adapted.The purpose of this paper is twofold; 1) test geometry limitations on indirect SLS whichwere originally developed for polymer SLS, and 2) catalog phenomena and limitations associated with indirect SLS (some of which may not be present in polymer SLS).2.

Materials and MethodsMaterialsThis work uses a blended alumina/nylon powder system as a testbed, which was adaptedfrom Deckers et al. [12]. A mixture of powders consisting of 78 wt.% alumina (Almatis A16 SG, d50=0.3µm) with 22 wt.% PA12 (ALM PA650 d50=58µm) was dry-mixed in a high- shear blender (Chulux QF-TB159008) for 10 minutes. The blended mixture was then sieved through a 250 µm mesh to produce the powders shown in Figure 2 and Figure 3.Figure 2: Schematic representation of morphology of the powder used in this work consisting of alumina and nylon (PA12)Figure 3: Scanning electron micrographs of blended alumina and nylon powders used in this work at a range of magnifications.689Methods: SLS Machine A custom SLS machine at The University of Texas at Austin, known as the Laser Additive Manufacturing Pilot System (LAMPS) was used for this study.

LAMPS has extensive sensing and control capabilities built into open-hardware and open-software architectures, which make it ideal for developing suitable process parameters for experimental materials [13].The parameters used for SLS were selected empirically with the help of visual and thermal imaging to ensure that degradation of the binder did not occur and that part curl was minimized [14]. The final parameters that were used are as follows:Laser power: 4 - 10 W Laser scan speed: 200 - 1000 mm/s Layer thickness: 100 µm Beam (hatch) spacing: 275 µm Spot size2: 1 𝑒2⁄ = 730 µm, full-width half-max (FWHM) = 580 µm Powder bed temperature3: 179 - 191°CMethods: Debinding and DensificationAfter the green parts were removed from the SLS machine, loose powder was removedwith a focused stream of dry air at a pressure of 680 kPa. Thermal debinding was performed in a tube furnace, with flowing dry air (~0.1 standard litre per minute) used to carry away the decomposition products.

Thermal debinding was performed in a tube furnace, with flowing dry air (~0.1 standard litre per minute) used to carry away the decomposition products. Each sample was placed in an alumina crucible on top of a thin layer of coarse-grained alumina (Goodfellow GF18024511, d50=45 µm) to avoid sticking of the part to the crucible. The crucibles were loosely covered to reduce direct contact with the flowing air. The parts were heated at 9°C/hour to 600°C, and then cooled to room-temperature naturally. After debinding, parts were transferred to a box furnace for final densification. Temperature was increased at a rate of 10°C/minute, and then held at 1600°C for 1 hour, before returning to room temperature at a rate of 10°C/minute.

Temperature was increased at a rate of 10°C/minute, and then held at 1600°C for 1 hour, before returning to room temperature at a rate of 10°C/minute. These heating and cooling rates were selected to minimize distortion of the part but were not optimized for overall processing time.Methods: Determining Geometry LimitationsAfter selecting a material system and processing route, a factorial-style approachthat was originally developed by Allison et al. was used to determine geometry limitations for samples containing through-thickness channels with two simple cross- sectional shapes: holes and slots (Figure 4 and Table 1). Laser power, and feature orientation were varied for each of these shapes.

Laser power, and feature orientation were varied for each of these shapes. There were two sub-methods for this procedure; 1) Qualitative: Visual inspection was used to determine if a feature has resolved, and 2) Quantitative: Pin gages were to measure dimensions of the holes and slots.The qualitative approach was used to assess if each feature (e.g., hole) resolved properly.In this work, a channel resolved if, by visual inspection, light passed through the channel. The quantitative approach used analysis of variance (ANOVA), with a confidence interval of 95%, to determine which factors significantly affected the outcome. The factors studied here were laser2 Measured via Ophir NanoScan 3 Monitored with a FLIR A6701 mid-wave infrared camera, which was used to control quartz lamps that heated the powder bed via PID control690power and feature orientation.

After performing the ANOVA analysis, post-hoc testing was used to quantify the degree to which the significant input factors affected the output. For example, if laser power significantly affected the measured hole diameter, by how much? This procedure is known as a multiple comparisons test or a Tukey-Kramer test [15]. An additional metric, the average mean difference, is included for clarity for some tests. The average mean difference determines the average amount that the feature of interest (e.g., hole diameter) is changed for a given type of input factor change (e.g., changing feature orientation from vertical to horizontal). The geometries shown in Figure 4 were produced in duplicate for each setting. So, for example, the part labeled “Large Holes” was produced twice vertically and twice horizontally for each laser power setting.Figure 4: Hole and slot channels.

So, for example, the part labeled “Large Holes” was produced twice vertically and twice horizontally for each laser power setting.Figure 4: Hole and slot channels. Hole diameters ranged from 0.8 mm - 4.7 mm with depths of 5 mm - 25 mm Slot widths ranged from 1 mm - 3.4 mm with depths of 5 mm- 25 mm. Slot height was held constant at 3.5 mm.Table 1: Features and input factors for holes and slotsFeatureFeature Range(s) & Part Specs.No. of Replicates (per feature per setting)Factor(s) & Setting(s) Scan Speed [mm/s]Laser Power [W]OrientationRasterHoleSlot0.8 - 4.7 mm hole diameter1.56 - 3.45 mm hydraulic diameter224.7, 6.5, 8.5, 9.1, 11.7, 144.7, 6.5, 8.510001000Horizontal, VerticalHorizontal, Vertical275 µm hatch spacing 275 µm hatch spacingThe quantitative comparisons presented for this study were measured using pin gagesinserted into the holes or slots.

Holes and slots measured using this method are characterized by their hydraulic diameters, Dhyd,hole or Dhyd,slot, so that comparisons can be made between them:𝐷ℎ𝑦𝑑,ℎ𝑜𝑙𝑒 = 𝐷ℎ𝑜𝑙𝑒𝐷ℎ𝑦𝑑,𝑠𝑙𝑜𝑡 =4𝑊𝐻 2𝑊 + 2𝐻𝑊 = 𝑠𝑙𝑜𝑡 𝑤𝑖𝑑𝑡ℎ, 𝐻 = 𝑠𝑙𝑜𝑡 ℎ𝑒𝑖𝑔ℎ𝑡, 𝐷ℎ𝑜𝑙𝑒 = ℎ𝑜𝑙𝑒 𝑑𝑖𝑎𝑚𝑒𝑡𝑒𝑟691(1)(2)The ratio of the nominal hydraulic diameter to the depth is defined from𝑁𝑜𝑚𝑖𝑛𝑎𝑙𝐻𝑦𝑑𝑟𝑎𝑢𝑙𝑖𝑐 𝐷𝑖𝑎𝑚𝑒𝑡𝑒𝑟 [𝑚𝑚] 𝑑𝑒𝑝𝑡ℎ [𝑚𝑚]≡𝐷 𝑑where nominal refers to the specified build dimension rather than the actual measured dimension.The first attempt to quantitatively measure the hole and slot hydraulic diameters wasadopted from Allison et al. Unfortunately, this approach, which used an optical flatbed scanner to measure holes sizes, was determined to be inadequate for the present work after some parts had already been densified. Therefore, a few comparisons with green parts are not possible because the final methodology for quantifying geometrical features had not been developed when these parts were densified.3.

Therefore, a few comparisons with green parts are not possible because the final methodology for quantifying geometrical features had not been developed when these parts were densified.3. ResultsA: Test of Geometry Limitations on Indirect SLS which were Developed for Polymer SLSQualitative Results There are two questions to be addressed in this section; I) are geometrical features in parts produced by indirect SLS less likely to resolve than in polymer parts, and II) Are the effects of processing parameters on geometrical limitations for SLS of polymer parts the same as for ceramic parts produced by indirect SLS.To address question (I), the “Large Holes” part from Figure 4 is used. Figure 5 shows anexample comparison between a nylon part produced by SLS by Alison et al. and a densified alumina part produced for this work. Figure 6 (A – B) addresses the differences in likelihood of resolution of holes in nylon parts and in alumina parts.

and a densified alumina part produced for this work. Figure 6 (A – B) addresses the differences in likelihood of resolution of holes in nylon parts and in alumina parts. Each cell contains a number that is the percentage of channels with that diameter and depth that resolved correctly. The boxes are shaded from dark-grey, which indicates that all of the parts resolved correctly, to white, which indicates that none of the parts resolved correctly.Figure 6 shows that for both materials, channels with large hole diameters and shallow depths are more likely to resolve properly than channels with small diameters and large depths. None of the channels in the ceramic part collapsed during debinding and densification, so the resolution tables are identical between the green and final parts.

None of the channels in the ceramic part collapsed during debinding and densification, so the resolution tables are identical between the green and final parts. Although the range of features studied is different between the material systems, within the overlapping region (e.g., hole depth 5-10 mm, hole diameter 0.8-2 mm), the nylon parts show a higher likelihood of resolving a feature in every instance.692(3)50mmFigure 5: Comparison of parts containing holes part in a nylon part from Allison et al. (Left) to holes in a densified alumina part from this work (Right)A) Holes in Nylon PartsB) Holes in Alumina PartsC) Slots in Alumina PartsD)Figure 6: Percentage of parts (indicated by number in each cell of the table) where features resolved for holes in nylon parts (A), reproduced from Allison et al. [9], holes in alumina (B), and slots in alumina (C).

[9], holes in alumina (B), and slots in alumina (C). Comparison of the results for all of the material systems (D).Question (II) is addressed by comparing the results for holes and slots in aluminaparts, which is shown in Figure 6 B and C, respectively. It is apparent that slots resolve to smaller diameters and larger depths than holes. For example, 50% of holes resolved for a 3 mm hydraulic diameter and 15 mm depth, but 100% of the slots resolved with the same hydraulic diameter and depth. This trend, that slots are more likely to resolve than holes, was also observed by Allison et al. for nylon parts, which indicates that the qualitative trends in feature resolution using indirect SLS of ceramics are similar to polymer SLS.693In summary, it is likely that the qualitative method developed for polymer SLS can also be applied to indirect SLS of alumina.

For most features, the likelihood of resolution is the same order of magnitude for both systems, and the qualitative trends are similar.Quantitative ResultsThis section will address similar questions to the qualitative section above, but now froma quantitative perspective. To address question (I), the accuracy relative to the specified dimension of holes and slots in alumina is compared to holes in nylon of similar specified dimension. As will be shown later, in some cases, the error in hydraulic diameters of the ceramic parts in their green state rivaled that of the nylon parts. However, shrinkage during debinding and densification increased this error significantly. The results from the densified alumina parts are presented in this section.Figure 7 shows that geometries produced by SLS in nylon are ~5 – 10× more accuratethan when produced by indirect SLS in alumina. In both material systems, slots are more accurate than holes on average.

In both material systems, slots are more accurate than holes on average. However, the deviation from specified dimensions is much larger in the alumina system (~20 - 80%) than in the nylon system (2 - 22%) for both holes and slots. It is likely that the larger error of the alumina system can be attributed to the heterogeneity of both agglomerate size and the distribution of materials in the powders used for indirect SLS.A)B)Figure 7: Average error and standard deviation of hydraulic diameter for slots and holes for densified alumina and nylon [Taken from 9]over all features and settings (A). Comparison between holes and slots for densified alumina parts and nylon parts (B).Question (II) is addressed by comparing a subset of similarly sized nylon and aluminachannels (Figure 8). The accuracy of the hole or slot is assessed by comparing the specified diameter to the diameter of that channel measured using a pin gauge. For holes in nylon parts, Allison et al.

The accuracy of the hole or slot is assessed by comparing the specified diameter to the diameter of that channel measured using a pin gauge. For holes in nylon parts, Allison et al. observed a decrease in the relative error in the diameter with decreasing hole depth. For slots in nylon parts, they found the accuracy to be relatively insensitive to the depth of the slot. In contrast, the relative errors for holes and slots in alumina parts produced for this study exhibit such large variances that they seem to be insensitive to the depth for the range tested. The data confirm that there are fundamental differences in the factors that limit the geometry for694indirect SLS compared to conventional SLS and that the quantitative trends observed in nylon parts are not replicated in indirect SLS.Figure 8: Left) Hole accuracy as a function of hole depth for alumina/nylon in green state and nylon-only holes part from Allison et al.[10].

Right) Slot accuracy as a function of slot depth.In summary, the data indicate that, the ability to produce a hole or slot that penetrates through the thickness of the part are similar for the two material systems. However, where the material systems differ significantly is in the measured accuracy of the holes and slots. The measured accuracy of these features is significantly poorer in alumina parts produced by indirect SLS compared to SLS of polymers.B: Catalog of Phenomena and Limitations Associated with Indirect SLSThe goal of this section is document phenomena that influence feature accuracy and that are specific to the process of producing alumina channels using indirect SLS. These phenomena were investigated via the quantitative procedure described earlier; ANOVA and post-hoc testing for each hole and slot at each depth was performed. Full results are presented in the appendix, but the pertinent information is included here.

Full results are presented in the appendix, but the pertinent information is included here. This section is separated into two parts; A) comparison of holes vs. slots, and B) phenomena associated with debinding and densification.Holes vs Slots To compare the phenomena associated with holes and slots, the relative error inhydraulic diameter versus the D/d ratio (Figure 9 A-B) is used in conjunction with the average mean differences (Figure 9 C). Mean differences are the relative changes in output, e.g., hole diameter, between each input settings whereas, average mean differences are the average of all mean differences for a given type of feature. Mean difference therefore can be used to explain how a particular feature (e.g., hole with depth of 10 mm and diameter of 2 mm) would respond to a given change in laser power. On the other hand, the average mean difference can be used to explain how a given change in laser power affects holes of all sizes.

On the other hand, the average mean difference can be used to explain how a given change in laser power affects holes of all sizes. The only input factor which can be directly compared is feature orientation since this parameter was systematically varied in both this study and the study by Allison et al. All of the input factors for both the Allison et al. work and this work are provided in Figure 9 C because their relative magnitudes are important.

All of the input factors for both the Allison et al. work and this work are provided in Figure 9 C because their relative magnitudes are important. The build parameters that were varied include: 1) the material (for nylon parts only) used for the build was Nylon 12, Nylon 12 reinforced with glass beads, and fire695retardant Nylon 11, 2) the feature orientation was either with the long axis of channel parallel or perpendicular to the laser beam 3) the location on the build platform (for the nylon parts only), and 4) the laser power used to produce the part (for the alumina samples only).Figure 9 A-B show the effects of specific build parameters on the relative error indiameter versus the D/d ratio. The color and shape of the data points specify the input factor (e.g., laser power). Figure 9 A does not show a significant influence of feature orientation on the relative error in hole diameter.

The color and shape of the data points specify the input factor (e.g., laser power). Figure 9 A does not show a significant influence of feature orientation on the relative error in hole diameter. In contrast, the relative error in hole diameter does appear to increase with laser power. However, a careful analysis using the full ANOVA results presented in the appendix showed that laser power is not a significant factor for all holes and hole depths. In contrast to the errors in hole dimensions, Figure 9 B suggests that the errors in the slot dimensions are somewhat agnostic to input factors, especially orientation; similar results were observed previously in nylon-only parts. [10], [11].A)B)C)Figure 9: Relative error in diameter versus the D/d ratio for holes (A), and slots (B). In (A), the solid blue line is for nylon holes from Allison et al.[10]. Scatter points are for densified alumina parts produced for this study.

In (A), the solid blue line is for nylon holes from Allison et al.[10]. Scatter points are for densified alumina parts produced for this study. Build variables for alumina parts are laser power and feature orientation. (C) Shows a comparison of average mean differences between nylon [11] and densified ceramic geometries. Percent significant indicates how often a given factor was significant. For example, orientation was a significant factor in alumina slots for 4 of the 12 instances studied.The average mean differences due to changing input factors on each feature are presentedin Figure 9. Interestingly, the average mean differences on hydraulic diameter are of the same order of magnitude between material systems. This means, for example, that the difference between slots built in horizontal and vertical orienations is similar between material systems (Δ=0.17mm for alumina, Δ=0.09mm for nylon).

This means, for example, that the difference between slots built in horizontal and vertical orienations is similar between material systems (Δ=0.17mm for alumina, Δ=0.09mm for nylon). Perhaps the most interesting conclusion from this figure is that holes and slots respond differently to input factors, which is seen in both material systems. This indicates that the differences between holes and slots results from the SLS process, rather than from the material systems themselves. The data for slots in alumina parts suggest only a weak dependence on laser power. The negative value in the orientation column indicates that slots built in the vertical orientation tend to be larger than slots built in the horizontal orientation. Note that this is not the case for slots in nylon parts, or for holes in either material system.

Note that this is not the case for slots in nylon parts, or for holes in either material system. This table should be used in conjunction with the full696ANOVA results in the appendix, which shows that orientation is only a significant factor for a minority of combinations of slot width and depth (33%).In summary, this section shows that holes and slots in alumina parts respond differently to input factors, but that this was also the case for holes and slots in nylon parts. Therefore, the phenomena cataloged here are not unique to alumina parts produced with indirect SLS, but are driven by the underlying physics of the SLS process.Debinding and DensificationThere is some indication from the results for holes presented in Figure 10 A that tuningthe process parameters for the indirect SLS system can produce green parts with similar accuracy to their nylon-only counterparts. However, it is possible that the accuracy can be further degraded during debinding and sintering.

However, it is possible that the accuracy can be further degraded during debinding and sintering. Deformation during debinding and densification has been well documented during conventional processing of ceramics (See Rahaman [16]). These phenomena are no different for indirect SLS, aside from the fact that slightly more binder is used in indirect SLS than traditional ceramics processes, such as powder injection molding (PIM). However, Figure 10 E shows that the errors are reduced in parts after densification. The standard deviation of error in hole diameter (over all hole sizes and input factors) for holes in the green state is 11.4%, while it is 7.3% for holes in densified parts. The mechanisms for this reduction in error will be the focus of a later study, but preliminary results are presented here.Figure 10 (C-D) shows that there is a slight decrease in the relative shrinkage of the hole with increasing hole diameter, but no statistical difference with feature orientation.

The plots also show that the relative shrinkage decreases with increasing laser power. Higher laser powers may result in larger melt zones that facilitate particle rearrangement during the sintering process. It is likely that this dependency is affected by the heterogeneous nature of the mixed alumina/nylon system. Further testing needs to be performed to confirm the cause of this phenomena.697A)B)C)D)E)Figure 10: Relative error in diameter versus the D/d ratio for holes in green state (A), and densified state (B). Relative shrinkage in diameter for holes varying orientation (C), and with varying laser power (D). Comparison of the average mean differences between green and final holes (E).ConclusionsWe have shown that the previously proposed geometry limitations for polymer SLS providea starting place for the design and manufacture of ceramic geometries using indirect SLS.

For example, green parts produced by indirect SLS show some similar dependencies on process input factors. However, the dimensional errors in alumina parts are considerably larger than those measured previously in nylon parts produced by SLS. In some cases, it is possible to improve dimensional accuracy of green parts to a level close to that of polymer parts by specifically tuning build parameters for indirect SLS. Even in these cases, binder removal and densification was shown to result in further degradation in the dimensional accuracy of alumina parts. Some trends in dimensional errors observed in direct SLS of nylon also occur in indirect SLS of alumina; for example, the accuracy of holes is more dependent on feature orientation than is the accuracy of slots.

There are also process-specific dependencies which warrant further research such as the dependence of part shrinkage on the laser power used to produce the green part.698AcknowledgementsThis work was supported by ExxonMobil through its membership in TheUniversity of Texas at Austin Energy Institute.Appendix A: Pictures of PartsFigure 11: Holes and slots in alumina/nylon in the green state. From top left, clockwise: A) Holes, laser power = 9.1 W, vertical feature orientation, B) Holes, laser power = 9.1 W, horizontal feature orientation, C) Slots, laser power = 8.5 W, vertical feature orientation, D) Slots, laser power = 8.5 W, horizontal feature orientationAppendix B: Full ANOVA ResultsThe following tables show significant effects and the mean differences associated witheach effect. The numbers inside of the parentheses show, on average, how much that factor affected hole diameter (in mm).

The numbers inside of the parentheses show, on average, how much that factor affected hole diameter (in mm). For example, for a hole diameter of 3.1 mm and depth of 5 mm, each change in power level decreased hole diameter by ~0.37 mm. Cells listed as N/A are instances which were not studied because their resolution likelihood was so small.Table 2: ANOVA and mean differences for holes in alumina-nylon in the green state. Letters indicate which factors were significant; P = Power; O = Orientation, P*O = interaction.Significant effects, (mean differences [mm])Hole Depth [mm] → Hole Diameter [mm] ↓ 1.5 2.3 3.13.94.75N/A N/A P(0.37) P(0.51), O(0.43), P*O P(0.74), O(0.36), P*O10N/A N/A N/AN/AP(0.71), O(0.55), P*O15N/A N/A N/AP(0.84), O(0.62)P(0.65), O(0.36)20N/A N/A N/AN/AP(0.78), O(0.55), P*O25N/A N/A N/AN/AP(0.72), O(0.54)699Table 3: ANOVA and mean differences for holes in a alumina after firing.

ometry Processing for SLSIHIP Ravi Venkataramani, Suman Das, Joseph J. Beaman Laboratory for Freefonn Fabrication University of Texas at AustinAbstractSLSIHIP is a new net shape manufacturing method that combines the strengths of direct selective laser sintering and hot isostatic pressing. Direct selective laser sintering is a rapid manufacturing technique that can produce high density metal parts of complex geornetry with an integral, gas iinpenneable skin. These parts can then be directly post-processed by containerless HIP. Sophisticated processing of the part geometry is required to facilitate the desired results from SLSIHIP. This paper presents geometry processing algorithms being developed for in-situ canning of SLSIHIP components.

This paper presents geometry processing algorithms being developed for in-situ canning of SLSIHIP components. This research is funded by DARPAIONR contract NO00 14-95-C-0 139 titled "Low Cost Metal Processing Using SLSIHIP".INTRODUCTIONSelective laser sintering (SLS) is a layered manufacturing technique that can produce freeform three-dimensional objects directly from their CAD models without part specific tooling or human intervention. Parts are built by selectively fusing layers of a powder material using a scanning laser beam. Details on this process are described el~ewhere'.~. The next generation of selective laser sintering i.e. direct fabrication of functional metal and cermet components and tooling is under development at the University of exa as^.^. To produce full density rnetal parts having complex geometry, a novel net shape manufacturing technique called S L S I H I P ~ ~ ~ has been developed at the University of Texas.

To produce full density rnetal parts having complex geometry, a novel net shape manufacturing technique called S L S I H I P ~ ~ ~ has been developed at the University of Texas. The idea is to consolidate the interior of a component to 80% or higher density and to fabricate an integral gas impermeable skin or "can" at the part boundary in-situ. Sophisticated geometry processing is required to obtain the desired results. It is proposed to develop parainetric representations of the part contours at the slice level. These parainetric curves will then be offset in a suitable direction, to grow or shrink the curves, to generate skins. This parametric representation will also enable local shape change to compensate for any changes in dimension or shape that might occur after HIP post-processing for a specific part. These parametric representations for each layer are then used as the contours for SLS processing.

These parametric representations for each layer are then used as the contours for SLS processing. The SLS processed part can then be directly post-processed by containerless HIP to full density. The optimal thickness of the skin will depend on the HIP deformation model and the tolerance requirements. A final machining step will result in a part having the desired geometry, mechanical properties and tolerance.BACKGROUNDIn the SLSIHIP process (Figure l), the component is produced by selectively consolidating a metal powder with a laser beam layer by layer. While producing each layer, a gas impermeable high-density skin (> 98% density) is formed at the boundariesof the part. The interior of the part is laser processed to a high density typically exceeding 80%. Thus, the part is shaped and canned in-situ. The encapsulated part is evacuated, sealed and post-processed by containerless HIP to full density.

Thus, the part is shaped and canned in-situ. The encapsulated part is evacuated, sealed and post-processed by containerless HIP to full density. A final machining step may be applied if necessary.Laser beam scans high density skin for currentI / cross-sectionILaser beam scans interior of current crpss-sectionIn-situ canningDirect SLSFinal MachiningHot Isostatic PressingFigure 1 The SLSIHIP ProcessOffset curves have been used in industry for a range of applications some of tolerance analysis and CAD in the which are inold design, NC path planning, automobile industry. An offset curve is a curve parallel to the original curve at some offset distance. Offsets can be generated for implicit curves as well as freeform curves. Offset techniques have been well established for a series of parametric curves like Herrnite curves7, ~ - s ~ l i n e s ~ , NURBS~, Pythagorean Hodograph curves"' etc.

Offset techniques have been well established for a series of parametric curves like Herrnite curves7, ~ - s ~ l i n e s ~ , NURBS~, Pythagorean Hodograph curves"' etc. It is proposed to use B-splines and NURBS for the considered application.B-splines have been chosen because they are used widely in industry and have been well documented. However they cannot represent regular or implicit curves like conics precisely. NURBS, being rational, are more versatile in representing both freeform and regular curves precisely.

However they cannot represent regular or implicit curves like conics precisely. NURBS, being rational, are more versatile in representing both freeform and regular curves precisely. One of the drawbacks of using NURBS is that the offsets generated will not be rational requiring the use of special polynomial parametric curves like Pythagorean Hodograph curves, which surmount this problem.GOALSThis research proposes to address the following issues:Generation of skins for each slice contour data, modification of part geometry based on part-specific HIP deformation models, development of intelligent scan patterns to produce an integral, gas impermeable skin by SLS, and to incorporate the above in intelligent process control.PROCEDUREThe sliced CAD data file of the component considered gives the contours along each layer. A B-spline or a NURB is fit to the slice data points. Details on the formulation of B-splines and NURBS can be found e l ~ e w h e r e " ~ ' ~.

A B-spline or a NURB is fit to the slice data points. Details on the formulation of B-splines and NURBS can be found e l ~ e w h e r e " ~ ' ~ . This parametric curve is then used to generate the offset. Depending on the relative location of the part contours at the slice level it inay be required to grow or shrink the curve, thus denoting a positive or negative offset respectively.Offset ProcedurmeLet (x(t),v(t)) be a point on a parametric curve C. Then, the curve C, offset a distance k dhas points (x,(t),y,,(t)) that satisfy the equations1'where ( (x(t), y(t)) is the tangent vector at (x(t), y(t)) . This is illustrated in Figure 2.Figure 2 Offset ProcedureSome of these offset curves may have self-intersections forming loops as shown in Figure 3.

This is illustrated in Figure 2.Figure 2 Offset ProcedureSome of these offset curves may have self-intersections forming loops as shown in Figure 3. These loops need to be removed to get an intersection-free contour, giving the offset.Loop RemovalIn order to remove loop that may occur in the offset generated the following sequence of operations may be performed1) Analyze the offset for self-intersectionsI3. Self-intersections define multiple closed ~ o o ~ s ' ~ .2) If the direction of rotation of such a loop is inconsistent with the direction of rotation of the original contour, it is removed.3) For loops whose direction of rotation is consistent, check the span of the remaining part of the offset.4) If the span of the remainder is greater, discard the loop.

5 ) If the span of the loop is greater, keep the loop and discard the rest of the offset.The result of the given set of operations on an offset curve with self-intersections is shown in Figure 4.0 OffsetFigure 3 Offset with loopsoop RernovalFigure 4 Loop Removal\ - \RESULTSTwo test cases were considered to validate the algorithm, one for a freeform contour and the other for a regular or implicit contour.Case I : Slice Contour o f a Stator VaneThis is one of the candidate components chosen to demonstrate the capabilities of SLS/HIP.Figure 5 Growing of the Stator VaneFigure 6 Shrinking of the Stator VaneCase 2 : Curve with ConvexitiesaaOffsetLoop RemovalFigure 7 Offset of an Implicit CurveThe polar form of the curve of Figure 7 is given bySUMMARYAn offset technique for B-splines was developed and tested for a series of curves. This technique will be validated by producing an SLSiHIP component.

This technique will be validated by producing an SLSiHIP component. A similar technique for NURBS should be developed to optimize contour specific processing. Part geometry re-engineering will be performed based on HIP deformation models.This research is supported by DARPAIONR contract #N00014-95-C-0 139 titled "Low Cost Metal Processing Using SLSIHIP".REFERENCESI Beainan, Joseph J. and Deckard, Carl R., Solid Freeform Fabrication and Selective Powder Sir2 tering, 1 51h N AMRC, North American Manufacturing Research Conference Proceedings, 1987, pp. 636-640. 2 Deckard, C.R., Ph.D. Dissertation, Department of Mechanical Engineering, The University of Texas at Austin, 1988. 3 Fuesting, et al, Development of Direct SLS Processing for Production of Cermet Composite Turbine Sealing Comporzerzts-Par-t I, Solid Freeforrn Fabrication Symposium 1996 Proceedings, pp. 39-46.

Glance at the recent additive manufacturing research and development in China 1Xiaodong Xing, 2Li Yang1Harbin Engineering University, Harbin, China 150012University of Louisville, Louisville, KY 40292Abstract This paper reviews some of the recent additive manufacturing research and development works in China. A considerable amount of AM research activities in China focuses on directed energy deposition processes, powder bed fusion processes and stereolithography, with much of the effect dedicated to system and application development. Although many of the recent results are not readily available from the literatures published in China, from the available information the areas of focus for research and development could be clearly seen. Despite some speculations, the AM research in China is vibrate and aggressive, with some areas at least several years ahead of the other countries.1.

Despite some speculations, the AM research in China is vibrate and aggressive, with some areas at least several years ahead of the other countries.1. Introduction China is among the earliest countries that started additive manufacturing (AM) research. Back in around 1990, several groups in China had started AM various AM research efforts, which include Tsinghua university (led by Dr. Yongnian Yan), Huazhong University of Science and Technology (led by Dr. Yungan Wang) and Xi’an Jiao Tong University (led by Dr. Binheng Lu). After over 20 years, the AM research in China has greatly expanded into a wide range of areas from aerospace, defense, automobile, biomedicine to appliance, tooling, micro/nano‐fabrication and art design.

Currently there exist over 10 large research groups and companies in China that have been extensively involved in AM research, which include Northwestern Polytechnical University, Beihang University, South China University of Technology, Nanjing University of Aeronautics and Astronautics, University of Science and Technology of China, Shanghai Jiao Tong University, Northwest Institute for Nonferrous Metal Research, China Academy of Engineering Physics and Beijing Longyuan beside the other three mentioned previously. Unlike some of the other countries such as U.S. and U.K., most of the research institutes in China also own their companies that sell their own systems and provide AM services. Therefore, one of the unique characteristics of AM research in China is that a significant portion of research efforts is dedicated to the system integration including both hardware and software.

Therefore, one of the unique characteristics of AM research in China is that a significant portion of research efforts is dedicated to the system integration including both hardware and software. On the other hand, the AM market in China is still largely focused on a few high value‐added applications, while the overall manufacturing market as well as the personal desktop printer market largely untapped. Likely due to the market demands, several AM processes including stereolithography (SLA), powder bed fusion (PBF) and directed energy deposition (DED) have seen quite considerable development in China in the recent years. For example, China is among the most advanced countries in the use of directed energy deposition technologies in the manufacturing of large aerospace components. On the other hand, some of these achievements are largely unfamiliarto the researchers outside of China.

On the other hand, some of these achievements are largely unfamiliarto the researchers outside of China. Therefore, in this paper, we attempt to review some of these works that have been published in various Chinese science and engineering journals. Although many of these publications did not offer sufficient technical details about certain works, overall some general observations can be made for the process/material development and applications. Also, as research areas such as bio‐printing and organ printing represent a rather unique branch of AM, these works were not within the scope of this review despite the fact that extensive research has been performed.2. Stereolithography (SLA) Multiple research groups in China started research works related to SLA back in early 1990s. Recently, more application oriented research works were performed by groups such as Xi’an Jiao Tong University by Dr. Binheng Lu and Dr. Dichen Li.

Recently, more application oriented research works were performed by groups such as Xi’an Jiao Tong University by Dr. Binheng Lu and Dr. Dichen Li. On the other hand, relatively limited works are focused on new photopolymer development [1,2].2.1 Equipment A lot of the research was focused on the development of new light source and the improvement of process accuracy. Several groups explored the use of non‐laser UV light source in SLA systems including mercury‐xenon UV lamp [3, 4, 5] and UV‐LED [6]. Mercury‐xenon UV lamp was used as a low‐cost substitute of the traditional UV laser sources, however suffers from low coherence, which requires additional optical manipulations [3, 4]. Jun et al. reported the development of a SLA system with LED as energy source, which could achieve a stable output power of over 30mW and a beam diameter of about 0.3mm [6]. Wu et al.

Jun et al. reported the development of a SLA system with LED as energy source, which could achieve a stable output power of over 30mW and a beam diameter of about 0.3mm [6]. Wu et al. suggested that the LED UV SLA system could result in significant cost saving as well as an energy saving of over 99%, although currently it also suffers the disadvantage of lower energy density and coarser beam size [4]. In the effort of improving control and fabrication accuracy, various techniques were studied by multiple groups, including the laser based resin position measurement [7], surface‐constrained recoating [8], and resin compensation [9], which were implemented in different commercial systems.

Various commercial SLA systems are available from companies such as Shanxi Hengtong and Beijing Yinghua, which are capable of fabricating photopolymer and ceramic parts with a layer thickness of 0.04mm and accuracy of 0.08mm [10], although layer thickness of 0.01‐ 0.02mm was also reported for research systems [8]. In addition, Shanxi Hengtong also released their digital light projection (DLP) based SLA system. Fig.1 shows some of the SLA systems developed by the same company.a. SCPS350B (DLP based)b. SPS600c. SPS250M (medical grade)d. SPS350C (ceramic)Fig.1 SLA systems from Shanxi Hengtong2.2 Process development A considerable portion of the researches related to the SLA process in China focuses on the accuracy improvement, which approach the problem from various aspects such as scanning strategy [11‐15], support generation [16] and residual stress reduction [17].

As many research groups develop their own SLA systems and therefore have full control of all process parameters, some interesting process strategies were investigated. For example, Lei et al. combined the double‐scanning curing strategy with the scanning area sectioning in order to reduce warping of overhanging structures, which showed significant improvement compared to both the original double‐scanning strategy and the STAR‐WEAVE strategy [11]. The same group also suggested that the proper design of sectioning and section scanning strategy based on the curing characteristic of the photopolymer could along potentially result in significant reduction of part warping and improvement of part quality [17]. In another work, through experimental based process optimization, Xiang et al. reported the successful fabrication of micro‐gears with minimum feature size of about 70μm using standard SLA system [18]. Hong et al.

reported the successful fabrication of micro‐gears with minimum feature size of about 70μm using standard SLA system [18]. Hong et al. developed an algorithm that recognizes different types of overhanging structure and generates support structures accordingly, which was implemented into the commercial systems by Shanxi Hengtong [16]. In addition, The group at Xi’an Jiao Tong University also performed extensive research works on the material and process development of specific systems, such as the investigation of time‐dependent curing profile evolution for ceramic particle loaded photopolymer and regular photopolymer with mercury‐xenon UV source utilizing both experimentation and molecular dynamics simulation using Monte Carlo method [15, 19].Due to the conflict of commercial interest between different research institutes, much of the process/material development research works published in China represent independent but often overlapping efforts with same problems.

In general, due to the focus on commercialization as well as the relatively mature level of technological development, there exist limited recent research works in China that focus on the process/material development of SLA.2.3 Applications With the rapid development of biomedical and aerospace industries in China, various studies have investigated the use of SLA for these applications, especially for those with stringent requirements with geometrical accuracies. Xi’an Jiao Tong University has worked extensively on the fabrication of wind tunnel and ceramic turbine blades using SLA as either direct or indirect methods, as shown in Fig.2 [20‐23]. Due to the limitation of material strength, the application of SLA in aerospace areas primarily focuses on the fabrication of molds and patterns that are subsequently used for investment casting.

Due to the limitation of material strength, the application of SLA in aerospace areas primarily focuses on the fabrication of molds and patterns that are subsequently used for investment casting. The parts fabricated by SLA could either be used directly as pattern for casting or as shells that can be reinforced by metal‐resin composites to serve as the mold for wax pattern casting [24‐26]. This type of processes was studied in various works in details using either finite element simulation or experimentation, which reported the resulting alleviation of thermal residual stress due to the casting process and the elimination of thermal cracks [25, 26].a. Wind tunnel [22]b. Turbine blade [23]Fig.2 SLA applications in aerospaceIn biomedical areas, extensive studies have been performed in various applications utilizing SLA, which is led by the group in Xi’an Jiao Tong University.

These works primarily focused on the fabrication of scaffolds for soft tissue regeneration or for hard tissue implantation [27‐34]. Zhuang et al. developed a three‐step process for the fabrication of PEG (polyethylene glycol)/ PLA (polylactide)/β‐TCP (tricalcium phosphate) composite bone/cartilage joint scaffold shown in Fig.3 [30, 31]. The β‐TCP structure was realized from the pattern fabricated by SLA, the PLA anchor structure was fabricated by weaving PLA fibers that mimic the ligaments, and the PEG cartilage was fabricated directly by SLA. This hybrid manufacturing method was also applied to fabricate ligament‐bone composite scaffold [33] and magnesium‐ceramic composite bone scaffold [34], which generally resulted in improved biological and biomechanical responses in both in‐vivo and clinical experiments. For hard tissue prostheses, using SLA parts as investment casting pattern, Yaxiong et al. have implemented custom AM titanium implants in more than 100 clinical trials [28].

For hard tissue prostheses, using SLA parts as investment casting pattern, Yaxiong et al. have implemented custom AM titanium implants in more than 100 clinical trials [28]. Overall the adoption of AM in biomedical applications has been significantly more aggressive than most other countries including U.S., which can be expected to serve as a critical advantage for the development of this technology in the long term.a. Large joint scaffold [30]b. Ligament‐bone scaffold [31, 35]Fig.3 Composite scaffold via SLA3. Powder bed fusion (PBF) Various groups in China have been extensively involved in PBF related research, which include Tsinghua University, Northwest Institute for Nonferrous Metal Research, South China University of Technology, Huazhong University of Science and Technology and Nanjing University of Aeronautics and Astronautics.

Most of these research groups have developed their own PBF systems, such as the EBSM‐150 electron beam PBF from Tsinghua University, DiMetal‐ 240/280/100 laser melting PBF from South China University of Technology, RAP‐1 laser PBF fromNanjing University of Aeronautics and Astronautics, and various laser polymer PBF systems from Huazhong University of Science and Technology.3.1 Equipment 3.1.1 Electron beam melting systems Tsinghua University developed its first electron beam based PBF system named electron beam selective melting (EBSM) around 2004, which was among the first research groups that independently developed such type of systems [36]. The research group led by Dr. Feng Lin has performed extensive research with the EBSM system including powder spreading, powder preheat and scanning control.

The research group led by Dr. Feng Lin has performed extensive research with the EBSM system including powder spreading, powder preheat and scanning control. The prototype system utilizes a hopper‐roller powder spread mechanism as shown in Fig.4a [37], which was later replaced by a vibration based powder spread mechanism that could achieve minimum layer thickness of 0.1mm via closed‐loop powder weight control [38]. Recently, Chao et al. further developed a vibration based dual‐ powder spread system as shown in Fig.4b, which could realize the controlled mixing of two powders [39]. It was suggested that this new mechanism could realize more efficient fabrication of digital materials and functionally graded materials. In addition, the same group has also demonstrated the successful spread of titanium powder with very low flowability and spreadability that were non‐printable via traditional powder spread mechanism.

In addition, the same group has also demonstrated the successful spread of titanium powder with very low flowability and spreadability that were non‐printable via traditional powder spread mechanism. Furthermore, a tilted comb spreader was also designed for the EBSM system, which could reduce shear stress exerted on the fabricated parts during the powder spreading [39]. Beside this group, multiple other groups also performed extensive research on platform development [40, 41]. The selective electron beam (SEBM) S1 system developed by Northwest Institute for Nonferrous Metal Research was reported to have layer thickness of 50‐200μm, beam size 200μm, scanning speed of 8km/s, material melting speed 10‐100m/s and fabrication accuracy of 1mm [40]. Overall these systems still lack the level of system integration and process accuracy of the commercial electron beam system by Arcam.

Overall these systems still lack the level of system integration and process accuracy of the commercial electron beam system by Arcam. However, due to their open architecture hardware and software, these systems are currently sought after as research platforms for new material development with electron beam energy sources.a. Hopper‐roller systemb. Vibration based dual‐powder systemFig.4 Powder spread mechanism in EBSM3.1.2 Laser melting systems South China University of Technology was among the first groups in China to develop laser melting PBF systems, and some of their systems are shown in Fig.5. Led by Dr. Yongqiang Yang, this research group also performed extensive research on the material development using theirin‐house laser melting systems, which will be introduced in the following section [42, 43].

Huazhong University of Science and Technology also developed their own HRPM‐II laser melting PBF system, and performed extensive research works in the improvement of laser deflection and focusing mechanisms using various approaches [44‐46]. It was reported that the HRPM‐II system has a positioning repeatability of 30μm and scanning accuracy of 0.1mm/100mm [46].a. DiMetal‐240b. DiMetal‐280 Fig.5 DiMetal laser melting systemsc. DiMetal‐1003.1.3 Laser sintering systems Huzhong University of Science and Technology is among the earliest groups in China that developed polymer laser sintering PBF systems [47‐49]. Wuhan Binhu Mechanical & Electrical Co. Ltd, which is affiliated to the university, currently offers at least six different models of laser sintering systems, which include the largest HRPS‐VIII system with build envelop of 1400x1400x500mm, layer thickness of 0.08‐0.3mm and part accuracy of 0.2mm/200mm [50].

Recently, the Changsha based Farsoon also started to provide laser sintering solutions from equipment to material supply [51]. By collaborating with BASF, Farsoon is also developing new polymer materials suitable for laser sintering PBF, although no further details are currently available. Fig.6 shows some of the commercial laser sintering PBF systems developed in China.a. HRPS‐IIIb. Farsoon 402c. Lasercore‐7000d. MEM450AFig.6 Laser sintering PBF systems in China3.2 Process development For electron beam PBF, extensive works have been carried out in the investigation of relationships between scanning strategies and the microstructure and properties of the fabricated parts. Qi et al. developed a kinematic model for the powder bed, which was subsequently used for the design of vacuum system in the EBSM system and the optimization of preheat strategies [52]. In another work, He et al.

In another work, He et al. proposed an electron beam preheat strategy, which gradually increases the preheating electron beam current in order to realize preheat of Ti6Al4V powder with relatively low electrical conductivity [53]. In addition, it was also suggested that the pre‐sintering effect could help reduce the tendency of balling during the melting process [53]. Several studies investigated the effect of powder mixing on their processiability, and concluded that when the mixing ratio between the more spherical gas atomized powder and the irregular water atomized powder is around 40%:60%, the powder bed exhibits highest processibility with both preheating and melting [54, 55]. In a recent work, Guo et al. studied the relationship between the input energy density and the surface texture of the 316L stainless steel fabricated via EBSM, and proposed a two‐step process strategy that aimed to improve the surface finish of each newly fabricated layer [56].

The proposed method utilizes low energy density scanning to introduce partial melting in the fabricated layer, which is followed by a subsequent re‐melting with higher energy density. It was reported that this approach potentially allows for more sufficient melting pool evolution by forming large‐size melting pools, which resulted in a top surface finish of less than 8µm, as well as 99.96% density with microstructure dominated by refined columnar and equiaxial grains [56].For metal laser melting PBF, extensive studies have been carried out by various groups that focus on the process‐microstructure‐property relationships of the laser melting PBF systems using various materials including 316L stainless steel, CoCrMo alloys, 304 stainless steel and Inconel 718 alloy [57‐70]. Wu et al. developed a core‐shell process strategy for the rapid fabrication of 316L stainless steel parts with fully dense shell and porous cores.

Wu et al. developed a core‐shell process strategy for the rapid fabrication of 316L stainless steel parts with fully dense shell and porous cores. Using the laser laser power (85W), higher scanning speed and larger scanning spacing (700mm/s and 0.08mm, respectively) were used for the core fabrication, while lower scanning speed and smaller scanning spacing (300mm/s and 0.06mm, respectively) were used for the skin fabrication. It wasreported that this approach helped reducing the thermal distortion of the fabricated parts by about 75% and process time by about 25% [60]. Song et al. studied the effect of laser process parameters (power, scanning speed and spacing) on the mechanical properties of the CoCrMo alloy, and it was reported that the higher yield strength, ultimate strength as well as hardness of the as‐fabricated parts were superior compared to the ASTM F75 specifications, while the elongation was significantly lower [61]. Lu et al.

Lu et al. studied the in‐situ alloying and fabrication of Fe‐ 8Ni‐0.5C alloy from pre‐mixed elemental powder [65]. The process parameters were optimized through single‐layer study, and a metal lattice part as shown in Fig.7a was successfully fabricated with 22A power, 45mm/s scanning speed, 0.07mm spacing and layer thickness of 0.18mm. Subsequent microscopy observed predominantly needle‐shaped martensite in the microstructure as well as some retained austenite as shown in Fig.7b‐c, while no significant element segregation was reported [65]. Zhang et al. compared the process parameters of a series of aluminum alloys including AlSi25, AlSi10mg, AlMg5 and AlMgSi0.5, and analyzed the mechanisms of internal porosity generations [66]. It was suggested that at least two porosity generation mechanism exist. When the input energy is insufficient, internal porosity could form due to the lack of fusion.

It was suggested that at least two porosity generation mechanism exist. When the input energy is insufficient, internal porosity could form due to the lack of fusion. On the other hand, when excessive input energy is used, the evaporation of magnesium contributes to the formation of small spherical porosities, which also result in the change of overall material compositions. The group at Nanjing University of Aeronautics and Astronautics led by Dr. Dongdong Gu has also performed extensive research with the process development for various materials including Inconel 718 and various metal matrix composite materials. Ying et al. studied the influence of input energy density to the solidification characteristics of the molten material [71]. It was found that at low input energy density, the microstructure of the IN718 parts fabricated by laser melting PBF exhibit large columnar γ grains with large cracks within these grains as well as small grains at the grain boundaries.

In addition, there also exists significant epitaph growth in the structure. As the input energy density increases, the IN718 microstructure start to exhibit more aligned columnar grains with less epitaph growth and refined grain size [71]. Meng et al. studied the fabrication of TiC reinforced Titanium matrix composite structure using laser melting PBF [67]. Using pre‐ mixed TiC‐Ti powder, the group investigated the effect of different processing parameters on the density and defects of the fabricated parts, and discussed the mechanism for the existence of an optimum input energy density. As shown in Fig.8, the microstructure of the TiC‐Ti composite exhibit homogeneous microstructure with nano‐size TiC phase dispersed within the Ti matrix until the energy density becomes excessive, at which point the microstructure becomes large dendrite with large grain sizes [67].

The same group also performed some very interesting studies with the process of non‐traditional materials and structures, such as the modeling of the melting pool dynamics for W‐Cu dual‐material system with tungsten particles remain solid during the melting process [68], as well as the use of foaming agent in the direct laser fabrication of 316L stainless steel with honeycomb porosities as shown in Fig.9 [69].1619a. Fabricated partb. Microstructure 1000xc. Microstructure 10000xFig.7 Fe‐Ni‐C parts by laser melting system via in‐situ alloying [65]a. 200J/mb. 267J/mc. 400J/md. 800J/mFig.8 Microstructure of TiC/C composite under different energy densities [67]a. Surfaceb. Cross‐sectionFig.9 316L stainless steel with honeycomb porosity fabricated by laser melting PBF [69]The research group in Huazhong University of Science and Technology led by Dr.

Surfaceb. Cross‐sectionFig.9 316L stainless steel with honeycomb porosity fabricated by laser melting PBF [69]The research group in Huazhong University of Science and Technology led by Dr. Yusheng Shi have performed systematic research with polymer laser sintering processes, while some of the other research groups focused more on non‐traditional sintering materials. Recently, Yan et al. studied the sintering phenomenon during the laser sintering process with several different materials, which found that the undesired secondary sintering between the processed region and the surrounding areas could be reduced by adding inorganic filler materials with high melting temperature [72]. It was also suggested that the secondary sintering is more significant for polymers with higher degree of crystallinity [73]. As it is widely known that current laser sintering process has considerable degree of uncertainty in process control. In an attempt to address this issue, Liu et al.

As it is widely known that current laser sintering process has considerable degree of uncertainty in process control. In an attempt to address this issue, Liu et al. explored the feasibility of implementing neural network based control algorithm for the process control optimization, although the authors did not find further works that verify its efficiency [74]. The same group also developed a hybrid process that combines laser sintering with hot isostatic pressing (HIP), in which the laser sintered parts wereused as the pattern for HIP mold, and the HIP was used as the primary shape generation process [75].

This approach overcomes some of the limitations of the laser sintering process in the manufacturing of metal parts, and was demonstrated to be capable of producing high strength AlSi316L parts with complex geometry as shown in Fig.10.Fig.10 Complext AlSi316L parts fabricated by SLS/HIP3.3 Applications Despite the extensive research, currently there exist relatively limited applications for metal PBF processes in China. Various groups have demonstrated the capability of PBF in the direct fabrication of metal structures and components in a range of applications including orthopedics [76, 77], surgical devices [78], complex mechanical joints [79], turbine engine plates [80] and emboss sculpture [81]. However, only a few areas that have high demand of high value‐added metal components have adopted metal PBF processes, such as aerospace and orthopedics.

However, only a few areas that have high demand of high value‐added metal components have adopted metal PBF processes, such as aerospace and orthopedics. On the other hand, the polymer and non‐metal PBF processes have been widely used in tooling industries to fabricate wax patterns, sand cores and even direct tooling for a range of industries including automobile and aerospace [82, 83]. Fig.11 shows a pair of casting mold for an engine block of 2000mmx1000mmx450mm in size fabricated by laser sintering process, with each half of the mold assembled from two pieces [83].Fig.11 Engine block mold by laser sintering, left is drag and right is cope4. Directed energy deposition (DED) One of the most rapidly advancing AM areas in China is the DED, which has started to receive considerable attentions worldwide.

Directed energy deposition (DED) One of the most rapidly advancing AM areas in China is the DED, which has started to receive considerable attentions worldwide. Started in late 1990s, various groups in China have worked extensively on this type of processes, with research almost exclusively focused on the process‐ microstructure‐property studies with various aerospace materials using laser based systems. Currently the two leading research institutes are Northwestern Polytechnical University and Beihang University, which have collaborated with defense agencies and independently demonstrated the fabrication of large titanium aerospace components using in‐house developed laser DED systems.4.1 Equipment There currently exist several commercial laser DED systems in China, which are mostly developed by owned Xi’an Bright Ltd that is owned by Northwestern Polytechnical University.

Until 2012, this group has developed various laser DED systems that use CO2 laser, Nd:YAG laser, fiber laser and diode laser as energy sources depending on their applications, and have sold multiple of these equipment to Chinese aerospace companies as well as GE China Research Center. These systems can achieve atmospheric control of O2<10ppm and possess real‐time closed‐loop feedback control systems based on melting pool temperature, melting pool size and layer height [84]. Fig.12 shows two of the systems developed by Bright. Recently, Beihang University also found a commercial company that starts to offer laser DED equipment [85].a. LSF‐Vb. LSF‐IVFig.12 Laser DED by Xi’an Bright [84, 86]4.2 Process development The research group in Northwestern Polytechnical University led by Dr. Weidong Huang focus extensively on the solidification process and microstructural control during the laser DED processes.

Weidong Huang focus extensively on the solidification process and microstructural control during the laser DED processes. As part of the State Key Laboratory of Solidification Processing, this research group has developed relatively comprehensive expertise on the fabrication of large metal components with controlled thermal stress and microstructure. Starting from 1998, the research group in Beihang University as part of the Engineering Research Center of Ministry of Education on Laser Direct Manufacturing for Large Metallic Component led by Dr.

Starting from 1998, the research group in Beihang University as part of the Engineering Research Center of Ministry of Education on Laser Direct Manufacturing for Large Metallic Component led by Dr. Huaming Wang have also performed extensive research works on the development of process parameters of various aerospace alloys including titanium alloys, Ni‐superalloys, Fe‐superalloys, high strength steels and intermetallics, as well as their post heat treatment processes for improved microstructural and performance control.Various literature works are available for the process development of various titanium alloys using laser DED, including TC2 (Grade 3), TC4 (Ti6Al4V), TC17 (Ti‐5Al‐2Sn‐2Zr‐4Mo‐4Cr) and TC18 (Ti‐5Al‐4.75Mo‐4.75V‐1Cr‐1Fe) [87‐101], as well as the in‐situ alloying study of Ti‐xAl‐yV (X<11, Y<10) [102‐104], TiB+TiC/TA15 and Ti‐6Al‐2Zr‐Mo‐V alloys [105‐107]. Ma et al.

Ma et al. developed a fractal scanning strategy, which helped to improve the temperature field homogeneity due to the self‐similar pattern of the fractal scanning paths [108]. Combined with the partial laser beam offset strategy, this scanning method was reported to significantly improve the part quality byreducing insufficient fusion and internal porosity. In another study, in order to avoid non‐ uniform thermal dissipation characteristics in different part of the scanning area in each layer, Ma et al. also developed a geometrical style recognition based automatic scanning path generation and process parameter control [109]. For regions A and B as shown in Fig.13a, the automated algorithm generates different process parameters that reduces various manufacturing defects and improve the fabrication quality and accuracy for the direct manufacturing of the C919 aircraft wing chord shown in Fig.13b.

Although no research literature was found by the authors, it was explicitly suggested that the fatigue and creep properties of the fabricated components are either comparable or better compared to the traditionally manufactured parts.a. Geometrical style recognitionb. aircraft wing chordFig.13 Geometry‐based process adjustment of laser DED [109]Extensive works have also been performed on the development of 17‐4PH stainless steel for component repair [110‐112]. It was reported that the microstructure of the 17‐4PH fabricated by laser DED exhibit two types of martensite as well as precipitated secondary phases, which evolves into fine lath martensite with coarsened secondary phases during tempering [110, 111]. The same group also employed cellular automata in the modeling of solidification and dendrite growth [112], as well as epitaph growth on Fe‐C single crystal substrate [113]. In the experimental investigation of the solidification behavior of laser DED process, Wang et al.

In the experimental investigation of the solidification behavior of laser DED process, Wang et al. utilized transparent butanedinitrile‐2.0% ethanol alloy and realized real‐time simulation of melting pool formation and non‐equilibrium solidification that mimics the real laser DED process [114]. It was found that the solidification and microstructure are affected by the melting pool morphology as well as the grain orientation of the substrate. There also exist interesting works from the other research groups. For example, Wang et al. explored the use of ultrasonic vibration in laser DED fabrication of BT20 (Ti‐6Al‐1.5Mo‐1.5V) alloy [115]. Using excitation frequency of 19.56kHz, it was found that the internal porosity and resulting grain size showed considerable improvement, however the deposition rate were also negatively affected. Pi et al.

Pi et al. studied the self‐stabilizing effect during the laser DED process that facilitates smooth finish, and proposed to employ process spacing at off‐focus distance in order to utilize this effect [116]. Wang et al. studied the adjustment of powder delivery rate in order to improve the accuracy and surface quality of the tilted thin‐wall structures [117].4.3 Applications Laser DED technologies have been adopted in various areas in China, including aerospace, biomedical, tooling, automobile and shipbuilding. One of the most widely reported applications is the direct fabrication of large aerospace components. Fig.14a shows a Ti6Al4V C919 aircraft wing chord with 450x350x3000mm size fabricated by Northwestern Technical University [118]. It was reported that the long‐term dimensions stability of this part is smaller than 1mm without any post heat treatment. In addition, both the quasi‐static and fatigue properties of the component are better than the wrought parts.

In addition, both the quasi‐static and fatigue properties of the component are better than the wrought parts. Fig.14b shows a dual‐alloy bearing case based on Ni‐superalloy, which was fabricated with two material compositions in order to better accommodate the thermal characteristics between different parts of the aircraft engines [118]. Beihang University fabricated various aerospace components including corner case, fitting for aircraft seats (shown in Fig.14c), pelvic joints and other secondary titanium structural components, which have been installed in various models [119, 120]. It was reported that the material utilization was improved by 5 times, and the fabrication time and cost were reduced by 67% and 50% respectively. Recently this group also successfully fabricated large TA15 titanium part for main structural components in aircraft with overall dimension over 1730x250x230mm as shown in Fig.14d.a. C919 wing chord [118]b. Bearing case [118]c. Seat fitting [119]d.

C919 wing chord [118]b. Bearing case [118]c. Seat fitting [119]d. Large aircraft structural comonent [119]Fig.14 Aircraft components fabricated by laser DEDNorthwestern Polytechncial University has also been collaborating with the Fourth Military Medical University to fabricate dental prostheses using laser DED. As shown in Fig.15, it wasexpected that functionally graded structures with metal‐hydroxylapatite graded material composition and varying porosity can be directly fabricated to accommodate both biological and mechanical requirements for the prostheses [121]. It was reported that this concept has been proved with multiple material combinations including stainless steel, CoMoCr alloys, titanium alloys and Ti/HA graded materials via preliminary clinical trial [121].Fig.15 Dental crown and mandibular based plate by laser DED [121]5.

Conclusions It was the impression of the authors that the AM research in China is highly aggregated in several major research institutes, which is partly contributed by the research resources that these institutes are capable of attracting. Almost all the research groups in China developed their own research systems, and some of these systems have been commercialized. In general, much of the equipment focused research between different groups exist significant overlap, which makes it difficult to grasp the most recent advancement of the technologies.Many of the AM research works focus on application development, especially with the relatively well‐developed AM processes such as stereolithography and powder bed fusion. On the other hand, several research groups in China have demonstrated extensive research expertise and achievement with both the fundamental process development and application development for the directed energy deposition processes.

With aggressive adoption by aerospace and biomedical industries, the application of AM may experience a more rapid progress compared to most other countries.There currently does not exist a research collaboration platform between U.S. and China. On the other hand, several mechanisms exist for the collaboration between Europe and China. There exist considerable potentials in collaborations between U.S. and China, which appear to have complementary expertise in many areas.Acknowledgement The authors are grateful of the information provided by Dr. Feng Lin and Wenjun Ge from Tsinghua University and Dr. Dichen Li from Xi’an Jiao Tong University.References [1] Yugang Duan, Xuerang Wang, Suqin Wang, Dichen Li, Bingheng Lu. Research on a Low Curl Photocurable Resin for Stereolithography. Journal of Xi’an JiaoTong University. 35(2001), 11: 1155‐1159.[2] Dandan Liu, Xin Wang, Tian Liu, Jie Yu, Wenyong Guo. Preparation of Stereolithography Resin. Engineering Plastics Application.

Performance of AISI316L complicated components manufactured by laser forming. Infrared and Laser Engineering. 39(2010), 3: 516‐520.[76] Huiping Tang, Guangyu Yang, Haiyan Liu, Nan Liu, Peng Jing, Zengfeng Li, Yuanping Huang. Study on Biomedical Porous Ti‐6Al‐4V Alloy Fabricated by Electron Beam Selective Melting. Rare Metal Materials and Engineering. 43(2014), Suppl.1: 127‐131.[77] Changhui Song, Yongqiang Yang, Manhui Zhang, Jiakuo Yu, Bo Yang, Di Wang. Redesign and Selective Laser Melting Manufacturing of Femoral Component Based on Digital 3D Technology. Optics and Precision Engineering. 22(2014), 8: 2117‐2124.[78] All‐digital rapid design and manufacture of metal customized surgical guide plate. Weihui Wu, Yongqiang Yang, Xingrong He, Huanwen Ding. Optics and Precision Engineering. 18(2010), 5: 1135‐1141.[79] Di Wang, Ruicheng Liu, Yongqiang Yang. Clearance Design and Process Optimization of Non‐ Assembly Mechanisms Fabricated by Selective Laser Melting.

ain Refinement of Freeform Fabricated Ti-6Al-4V Alloy Using Beam/Arc ModulationScott Mitznera, Stephen Liua, Marcia Domackb, Robert HafleybAbstractGrain refinement can significantly improve the mechanical properties of freeform-fabricated Ti- 6Al-4V alloy, promoting increased strength and enhanced isotropy compared with coarser grained material. Large β-grains can lead to a segregated microstructure, in regard to both α- phase morphology and α-lath orientation. Beam modulation, which has been used in conventional fusion welding to promote grain refinement, is explored in this study for use in additive manufacturing processes including electron beam freeform fabrication (EBF3) and gas- tungsten arc (GTA) deposition to alter solidification behavior and produce a refined microstructure. The dynamic molten pool size, induced by beam modulation causes rapid heat flow variance and results in a more competitive grain growth environment, reducing grain size.

The dynamic molten pool size, induced by beam modulation causes rapid heat flow variance and results in a more competitive grain growth environment, reducing grain size. Consequently, improved isotropy and strength can be achieved with relatively small adjustments to deposition parameters.a Colorado School of Mines, Golden, CO b NASA Langley Research Center, Hampton, VAKey words: Ti-6Al-4V alloy, grain refinement, electron beam, gas-tungsten arc, freeform fabricationIntroductionSolid freeform fabrication(SFF) is an additive manufacturing process in which material isadded to a substrate to build a near-net-shape part. Like other forms of manufacturing, SFF utilizes heat sources typically used in material joining processes such as gas tungsten arc or electron beam. Most metallic forms of SFF use layer-additive methods for deposition.

Most metallic forms of SFF use layer-additive methods for deposition. The process typically uses a computer aided design of two-dimensional slices of a three dimensional part to describe the layer-by-layer deposition of material. Feedstock material can consist of powder, sheet or wire that is melted and fused to the previously deposited layer. There are many forms of this type of manufacturing, including Electron Beam FreeForm Fabrication (EBF3), Laser Engineered Net Shaping (LENSTM), Direct Metal Deposition (DMD), and Laser Additive Manufacturing (LAM). These processes are not only used for fabrication of individual components but for repair and salvaging of other parts, or addition of structural details onto simple castings and forgings.

These processes are not only used for fabrication of individual components but for repair and salvaging of other parts, or addition of structural details onto simple castings and forgings. Several industries, such as medical and aerospace, benefit from the energy and cost savings afforded by additive manufacturing, primarily due to the use of costly alloys in these fields.SFF has many advantages over traditional methods of manufacturing, such as energysavings, decreased material waste, shorter design-to-final fabrication time, and increased material properties. Traditional machining and manufacturing methods require certain tolerances and geometries that must be incorporated into the design of the part. These processes can also require expensive consumables for removing material. SFF requires access to only the last536surface layer giving the method a large advantage for tight areas or hollow parts.

SFF requires access to only the last536surface layer giving the method a large advantage for tight areas or hollow parts. Casting requires molds, dies and equipment for manufacturing, while SFF requires only a substrate to build upon, decreasing time and cost. The buy-to-fly ratio (ratio of purchased raw material to material used) can be as low as 2:1, where half of the material is used for the final part. Traditional manufacturing methods often result in 14:1 – 20:1 buy-to-fly ratio [1]. Lastly, the deposited microstructure of SFF offers significantly improved mechanical properties, particularly in comparison with cast parts. The localized heating and cooling also restricts chemical segregation, which is frequently observed in thick parts produced in conventional casting.Many researchers have characterized the microstructure of Ti-6Al-4V material deposited by traditional freeform fabrication methods [2-8].

The typical microstructure consists of the two phases common in Ti-6Al-4V, body centered cubic (BCC) β phase and hexagonal closed packed (HCP) α phase. Using most types of power sources and deposition strategies, large epitaxial β grains are grown in the direction of heat flow during additive manufacturing of Ti-6Al-4V components [2, 8-13]. These large columnar grains grow from the substrate and often result in anisotropic properties [8, 14-16]. Deposition parameters, power source type, consumable attributes and build dimensions can all have an effect on the final microstructure. Beam and arc modulation are common methods of microstructural refinement in traditional fusion welding of Ti-6Al-4V for many reasons, including decreased hot cracking and increased fracture toughness [17-20].By means of grain refining techniques like beam or arc modulation, mechanicalproperties may also be dramatically improved.

Very limited research has been done on the potential advantages of modulation to SFF [21]. Multiple mechanisms have been proposed regarding the grain refinement of metals during modulated welding. Dendritic fragmentation is the commonly attributed mechanism for grain refinement in modulated weld fusion zones [20, 22]. As the arc current decreases, the weld pool shrinks and the solidification front progresses towards the center of the weld pool at a temporarily increased rate. The solidification front is sensitive to disruptions while advancing towards the center. As the current of the arc then begins to increase, the growing dendrites or entire grains may break from the solidification front and remain solid in the weld pool acting as heterogeneous nucleation sites [23, 24].

However, dendritic fragmentation has been eliminated from the possible mechanisms in Ti-6Al-4V alloy due to the cellular and planar nature of the solidification front, rather than the dendritic solidification commonly observed in aluminum and steel alloys [22]. Increased fluid flow is another mechanism suggested to account for grain refinement due to pulsation from the repeated disruption of arc force. The varied arc forces cause vibrations in the weld pool, constantly adjusting weld pool size [20]. A result from increasing fluid flow and pool agitation is the lowered thermal gradient of the solidification front due to increased mixing [25]. Therefore, the peak thermal gradient of the solidification front is constantly varying, causing a more competitive growth environment and resulting in a relatively large amount of grains growing in small increments.

In comparison, only selected grains with preferred orientations grow to a larger extent in conventional welding processes [20, 25, 26]. In this study, the potential of beam and arc modulation are explored to refine the grain size in Ti-6Al-4V deposits fabricated by Electron Beam Freeform Fabrication (EBF3) and Gas Tungsten Arc (GTA) deposition, respectively. For each process, a range of process parameters is examined and grain refinement evaluated through quantitative metallographic techniques.537Experimental ProcedureThe NASA LaRC Electron Beam FreeForm Fabrication (EBF3) machine consists of a 42kW electron beam welder with a six-axis motion control system and dual wire feeders for wire deposition. The entire 72 in x 24 in x 24 in build volume is contained in a high vacuum environment during deposition. The experimental matrix used to evaluate the effects of beam modulation is shown in Table 1.

The entire 72 in x 24 in x 24 in build volume is contained in a high vacuum environment during deposition. The experimental matrix used to evaluate the effects of beam modulation is shown in Table 1. Powder core tubular wire (PCTW) manufactured at the Colorado School of Mines was used in the EBF3 system to mitigate vacuum-induced aluminum loss previously observed in the EBF3 process and correct the net chemical composition of the deposit [1]. The height of deposit, length of deposit, wire feed rate, traverse speed, and accelerating voltage were held constant while beam current, frequency of modulation, and background current were varied to identify the most effective modulation waveform condition and frequency for maximum microstructural refinement. Specific values for the deposition parameters are proprietary, consequently these parameters are shown in a normalized fashion in order to convey the range explored.

Specific values for the deposition parameters are proprietary, consequently these parameters are shown in a normalized fashion in order to convey the range explored. Two waveform conditions were explored, as shown in Figure 1, each at two modulation frequencies along with a non-modulated baseline. The total heat input was equivalent for both modulation conditions, but slightly lower than the non- modulated build. Both waveform conditions utilize peak current for 50% of the time and utilize background current for the other 50% of the time.Table 1Normalized deposition parameters for modulation experiments on EBF3.DepositionHeight (in)Length (in)Beam Current (Normalized) HighModulation Freq.

(Normalized)Background Current (%)Beam Current (Normalized) LowFraction of Time at Peak Current (%)Heat Input (Normalized)10.55.00.690N/AN/AN/A12 30.5 0.55.0 5.00.85 1.000.5f 0.5f50 250.43 0..2550 500.90 0.904 50.5 0.55.0 5.00.85 1.001.0f 1.0f50 250.43 0.2550 500.90 0.90538Figure 1Modulation waveform schematics for use with EBF3: (a) 50% normalized background current; (b) 25% normalized background current.The Colorado School of Mines Gas-Tungsten Arc freeform fabrication machine uses a modified automated Miller Dynasty 350 GTA welding system with a torch enclosed in a box with controlled atmosphere. Figure 2 is a schematic of the machine. The railed deposition stage is connected to a shaft which passes through the wall of the chamber with a sealed pass-through. The torch and wire feed system is attached to another shaft for torch height manipulation during deposition of layers.

The torch and wire feed system is attached to another shaft for torch height manipulation during deposition of layers. For environmental control, a small rough vacuum pump is attached on one wall opposite the gas inlet, through which UHP argon is flowed at 40 l/hr to maintain atmospheric pressure in the chamber. The machine is capable of producing 8 in long x 2 in tall wall builds. Solid Ti-6Al-4V 1.2 mm diameter welding wire was deposited onto ground and cleaned CP Ti plate. Arc modulation waveform conditions are shown in Figure 3. In contrast to the EBF3 builds, background current remained at a constant 40% of peak current for the GTA builds, while the peak current time was varied between 50% and 15%. The heat input for both modulated and non-modulated builds were held constant.

(a) 50% background current time (b) 15% background current time.Upon completion of the wall builds, specimens were cross-sectioned along the Y-Z planeaccording to the coordinate system illustrated in Figure 4. The cross-sections were mounted in Bakelite and ground through 1200 grit SiC paper. The samples were then polished with 6 μm and 1 μm diamond solutions on green felt and nylon polishing pads, respectively. The samples were then placed on a vibratory polisher with 0.05 μm colloidal silica on a Lecloth pad for the final polish. For the macrographs only, the samples were etched with Kroll’s reagent to develop the microstructure for observation using an Olympus PMG3 microscope. Manual and semi- automated metallographic analysis methods were used to characterize four microstructural features identified in literature as having the largest effects on mechanical properties in Ti-6Al- 4V alloy [27, 28].

Ti-6Al-4V weld microstructures generally contain two phases, α and β, in a variety of morphologies including globular, basketweave and colony.Figure 4Coordinate system utilized in all experiments.540Two β phase features, grain size and volume fraction were analyzed in the steady-stateregion of each build. The steady –state region of the build was defined as the region of the deposit that several layers away from the baseplate. The β grain size was characterized using optical macrographs that included a number of grains for suitable statistics. The ImageJ software was used for the analysis using the “Circles” subroutine, which placed circles on the macrographs. A total of 318 circles, placed over the image in a random grid as shown in Figure 5, were used to analyze two macrographs.

A total of 318 circles, placed over the image in a random grid as shown in Figure 5, were used to analyze two macrographs. The intersection of the circle line and the β grain boundaries were then manually counted and an average intercept length calculated as the β grain size using Equation 1.Equation 1Calculation of β grain size:Figure 5The β grain size quantification method of circular intercept. Yellow dots represent intersections with β grain boundaries.The β phase volume fraction was determined using scanning electron microscope (SEM)back scatter imaging with a JEOL 7000F SEM and the ImageJ software. Five regions were analyzed as follows and as illustrated in Figure 6.

Five regions were analyzed as follows and as illustrated in Figure 6. The “Find Edges” subroutine was used to identify the regions of β phase, then the image was converted to binary contrast, and the volume fraction of β phase was determined using a macro program.541Figure 6Procedure of β phase fraction analysis: a) Initial backscatter image of an unetched sample showing the chemistry difference of the two phases; b) “Find Edges” command in ImageJ finds the β phase edges in the microstructure; c) Conversion to binary for calculation of β phase fraction.The α phase colony intercept length and lath width were also characterized in the steady-state region of each build. Figure 7 illustrates the methodology for calculating the average colony intercept length. One hundred random line segments of known length were placed on ten randomly chosen back scatter SEM images and each colony intercept was counted.

One hundred random line segments of known length were placed on ten randomly chosen back scatter SEM images and each colony intercept was counted. Using the total number of counts and total line length, an average colony intercept length was found using Equation 2.Equation 2Calculation of α colony intercept length:The α phase lath width was measured manually using a line segment of known lengthplaced perpendicular to the α phase colony direction, shown in Figure 8. Over 10 random micrographs, 300 laths were measured to find the average lath width for each build. Error analysis was performed according to ASTM E112 standard and the calculated relative accuracy is shown in Table 3. Values less than 10% indicate that these results are accurate. Error bars throughout this paper are represented by the calculated 95% confidence index.542Figure 7Colony intercept length is determined by the placement of random line segments of known length.

At the intersection of two colonies, notice the placement of a yellow dot. The number of the intercepts, yellow dots, per unit line length can be used to find average colony intercept length.Figure 8The method of calculating α lath width with perpendicular lines measuring width at the widest point of the lath.Initial mechanical property evaluation involved microhardness testing of the steady-state region of each build in all three planes: X-Z, Y-Z, and X-Y. Cross-sections were made through each section and prepared in the same manner as described earlier for microstructural characterization, through 0.05 μm colloidal silica, but these samples were not etched. Fifteen random Vickers hardness indentions in the steady-state region were completed with a Leco MHT Series 200 microhardness machine on each of the three planes for each build.

Fifteen random Vickers hardness indentions in the steady-state region were completed with a Leco MHT Series 200 microhardness machine on each of the three planes for each build. Indentions were performed with a 10 second dwell time and a 500 gram load.543Table 3Average relative accuracy of the quantification of microstructural features. Readings with 10% or less are considered accurate [29].Characterized Feature β Phase Volume Fraction α Phase Colony Intercept Length α Phase Lath Width β Phase Grain SizeAverage Relative Accuracy (%) 8.8 9.0 10.1 6.8Results and DiscussionBeam modulation during Electron Beam Freeform Fabrication:Electron beam modulation was implemented with a consistent 50% background timewaveform condition; however, the background current was varied at 50% and 25% of the peak current to investigate the effect of modulation condition.

It should be noted that the traditional, non-modulated builds had a slightly higher heat input than the beam modulated builds. All plots refer to the traditional non-modulated builds as a normalized frequency of modulation of 0.β Phase Refinement:With modulation, the traditional long columnar grains typical of electron beam freeform fabrication (EBF3) were modified to a much more equiaxed structure. Fresh grains were found to nucleate in the steady-state region of the build, where epitaxial growth normally predominates, as seen in Figure 9. The β grain size was drastically reduced for both modulation waveform conditions, for example, from 1164 μm to 734 μm with the 25% background current, as shown in Figure 10. Accordingly, volume fraction of β phase increased with modulation, from 22.7% to 35.0%, with the peak increase of β phase seen at the highest modulation frequency.

Accordingly, volume fraction of β phase increased with modulation, from 22.7% to 35.0%, with the peak increase of β phase seen at the highest modulation frequency. Utilizing backscatter micrographs shown in Figure 11, the variance of phase types can be observed.Figure 9β grain size refinement: (a) Non-modulated; (b) Modulation per deposition conditions for deposit 5 in Table 1.α Phase Refinement:The refinement of α phase was observed for both frequencies, with the most refinement atthe highest modulation frequency. Figure 12 shows representative micrographs for non- modulated and modulated builds. The microstructure for both the non-modulated and the modulated builds is primarily composed of fine colony α phase morphology with small areas of basketweave morphology. Figure 13 shows the refinement of the α phase due to modulation, as544indicated by the reduction in α colony intercept length from 3.2 μm to 2.2 μm.

Figure 13 shows the refinement of the α phase due to modulation, as544indicated by the reduction in α colony intercept length from 3.2 μm to 2.2 μm. The α lath width has also been refined by half, from 1.0 μm to 0.5 μm, also shown in Figure 13.Figure 10β grain size: (a) refinement using modulation for each waveform condition. The largest grain size refinement was observed at maximum modulation frequency. β phase volume fraction: (b) was increased from 22.7% to 35.0%.Microhardness of modulated EBF3 Builds:The mechanical response to the refinement in microstructure was measured with Vickers microhardness. Each plane, X-Z, Y-Z, and X-Y was analyzed to estimate mechanical properties in each direction. Figure 14 shows that hardness in all planes increased due to microstructure refinement, with the Y-Z plane showing an increase from 325 VHN to 356 VHN.

Figure 14 shows that hardness in all planes increased due to microstructure refinement, with the Y-Z plane showing an increase from 325 VHN to 356 VHN. Figure 14 also shows an increase in hardness for the X-Z and X-Y planes from 338 VHN to 353 VHN and 343 VHN to 356 VHN, respectively. The peak microstructural refinement occurred in the 25% background current waveform condition and peak hardness was consistent between the two modulation conditions. Consistent with the changes in α and β phases with frequency, hardness continued to increase over the range of frequencies examined.Figure 15 illustrates the improved material isotropy. The non-modulated builds,highlighted in the red boxes, yielded mixed hardness values for the three planes, while the moderate frequency for both modulation conditions yielded consistent increases in hardness over all planes.

Frequency seems to have the largest effect on hardness, as the observed hardness variation between the two waveform conditions was small.Figure 11β phase fraction: (a) Non-modulated; (b) Modulation per deposition conditions for deposit 5 in Table 1.545Figure 12Refinement achieved through modulation per deposition conditions for deposit 5 in Table 1: (a) baseline build with no modulation at lower magnification; (b) modulated build at lower magnification; (c) baseline build with no modulation at high magnification; (d) modulated build at high magnification.The Hall-Petch relationship is a method for determining how microstructure and mechanical properties impact each other using the relationship in Equation 3 [30]. Utilizing α lath width in the Y-Z plane as the characteristic distance, d, Figure 16 shows the good correlation between this relationship and the data.

Utilizing α lath width in the Y-Z plane as the characteristic distance, d, Figure 16 shows the good correlation between this relationship and the data. The regression coefficient R is 0.963 indicating the measured α width is a just assumption of the microstructural-microhardness relationship.Figure 13α colony intercept lengths (a) and α lath width (b) for both modulated conditions. Largest refinement and smallest lath spacing occurs for the deposition conditions for deposit 5 in Table 1.546Equation 3Figure 14Figure 15Hall-Petch relationship, where VHN is the hardness number and d is the characteristic distance (usually grain size):√Variation in microhardness with frequency measure in: (a) Y-Z plane, (b) X-Z plane, and (c) X-Y plane.Bar plot showing the hardness in each direction with non-modulated builds highlighted by the red boxes. Hardness has been increased for all directions, resulting in more isotropic values.

Hardness has been increased for all directions, resulting in more isotropic values. (a) Shows the 25% background current waveform condition (b) Shows the 50% background current waveform condition.547Figure 16Hall-Petch relationship utilizing α width. The regression coefficient of 0.96 indicates a strong linearity of the data.Arc modulation during Gas-Tungsten Arc freeform fabrication:Two modulation waveform conditions were utilized for the GTA deposits. In contrast to the EBF3 builds, which varied the background current while keeping the time at a constant 50% peak current time and 50% background current time, the GTA builds kept the background current constant at 40% of the peak current while varying the time at peak current. The two peak current times used were 50% peak current time and 15% peak current time to magnify the possible effects from modulation. Baseline builds were made with traditional non-modulated parameters.

Baseline builds were made with traditional non-modulated parameters. For all builds, parameters were varied in order to keep the total heat input constant to eliminate any possible effects from a change in heat input. All plots refer to the traditional non- modulated builds as a normalized frequency of modulation of 0.β Phase Refinement:Refinement of the large β grains found in non-modulated builds was successfullyaccomplished with both modulated waveform conditions. Figure 17 shows the morphology change due to the modulated arc. Similar to the EBF3 builds, epitaxial grain growth was still evident in the modulated builds; however new grains with much more equiaxed shape were found in the steady-state region. Figure 18 shows the change in β grain size as a function of frequency for both modulation conditions. In contrast to the EBF3 builds, the greatest degree of refinement occurred at an intermediate frequency of modulation.

In contrast to the EBF3 builds, the greatest degree of refinement occurred at an intermediate frequency of modulation. During the higher frequency builds the weld pool did not have sufficient time to change shape, likely due to the weakening effect of weld pool agitation, therefore the mechanism of grain refinement was inhibited. At peak refinement the β grain size decreased from 1450 μm to 980 μm, which is consistent with literature data [20]. A representative micrograph of the increase in β phase volume fraction is shown in Figure 19. The peak in β phase volume fraction, as shown in Figure 18, occurred at the same modulation waveform condition as the minimum in β grain size. The fraction of β phase increased from 14.7% to 24.6%.

The peak in β phase volume fraction, as shown in Figure 18, occurred at the same modulation waveform condition as the minimum in β grain size. The fraction of β phase increased from 14.7% to 24.6%. A decrease in β phase volume fraction was observed at the high frequencies, which is consistent with the evidence of loss of the grain refining mechanism observed for the β grain size measurements.548Figure 17β grain refinement due to arc modulation: (a) Non-modulated build macrostructure; (b) Modulated arc macrostructure as described for deposition 8 in Table 2.Figure 18β phase refinement due to arc modulation: (a) β grain size for both modulation conditions; (b) β phase volume fraction increase for both waveform conditions.α Phase Refinement:The effect of modulation on α phase morphology was characterized by α colony intercept length and α lath width. Representative backscatter electron micrographs are shown in Figure 20.

Representative backscatter electron micrographs are shown in Figure 20. Similar to the β phase refinement, a similar trend of peak refinement at an intermediate frequency followed by coarsening at higher frequencies was observed for α colony intercept length, as shown in Figure 20. The peak refinement occurred at the same conditions as for β phase refinement, with α colony intercept length reduced from 2.9 μm to 2.2 μm as compared with the non-modulated build. Figure 21 also shows that α lath width decreased with increasing frequency for both of the modulated conditions; however the behavior of α lath width did not show the microstructural coarsening in higher frequencies shown in the previous quantified features. Due to the theoretical maximum colony intercept length having dependence on β grain size, it is not surprising that the trend of colony refinement is similar to the β grain size trend.

Due to the theoretical maximum colony intercept length having dependence on β grain size, it is not surprising that the trend of colony refinement is similar to the β grain size trend. The α lath trend is unique, however, as it showed a constant refinement over all frequencies of modulation.549Figure 19Representative change in β phase fraction with arc modulation: (a) Non-modulated builds parameters; (b) Modulated arc build as described for deposit 8 in Table 2.Figure 20Refinement of α phase due to arc modulation: (a) Baseline build with no modulation at low magnification; (b) Modulated build at low magnification; (c) Baseline build with no modulation at high magnification; (d) Modulated build at high magnification.Microhardness of modulated GTA freeform fabrication builds:Microhardness was used to determine the effect of arc modulation and resulting grainrefinement on material properties.

Hardness on each of the planes, Y-Z, X-Z, and X-Y was measured as an indication of changes in strength and to examine anisotropy within the builds. The trend in hardness was similar for all planes and for both waveform conditions, as shown in Figure 22. All the arc-modulated builds displayed beneficial increases in hardness as compared with the baseline, non-modulated build. The trend in microhardness was similar to that observed for α lath width, characterized by a rapid change with arc modulation at the lowest frequency followed by a plateau in values for subsequent increase in frequency. The degree of anisotropy for each waveform condition is shown in Figure 23. A high degree of anisotropy is evident in the builds made with traditional non-modulated parameters, highlighted in red boxes in Figure 29. Arc modulation resulted in increased hardness for nearly all planes, and significantly550reduced anisotropy.

Arc modulation resulted in increased hardness for nearly all planes, and significantly550reduced anisotropy. Hall-Petch analysis was conducted for α lath width and microhardness for the Y-Z plane, and the results are plotted in Figure 24. The plot shows a good agreement between microhardness and α lath width, with a linear regression coefficient R2 of 0.90.Figure 21Trend in α phase refinement with frequency: (a) α colony intercept lengths for both modulated conditions; (b) α lath width for both modulation conditions.Figure 22Variation in microhardness with frequency measure in: (a) Y-Z plane; (b) X-Z plane; (c) X- Y plane.Figure 23Bar plot showing the hardness in each direction with non-modulated builds highlighted by the red boxes. Hardness was increased by arc modulation in all directions, resulting in a551more isotropic specimen.

Hardness was increased by arc modulation in all directions, resulting in a551more isotropic specimen. (a) 15% peak current time waveform condition, (b) 50% peak current time waveform condition.Figure 24Hall-Petch relationship of α lath width to VHN. The linear regression coefficient of 0.90 indicates a good linear relationship.ConclusionsMicrostructural refinement and improved mechanical isotropy in Ti-6Al-4V alloy builds was successfully demonstrated with changes to deposition parameters for both the EBF3 system at NASA LaRC and the GTA freeform fabrication machine at the Colorado School of Mines. The main conclusions drawn from this work are the following:1. Deposition using the EBF3 system showed a trend of continuously increasing grain refinement with beam modulation frequency, with peak refinement at the maximum frequency examined. β grain size decreased from 1160 μm to 730 μm and β phase volume fraction increased from 22.7% to 35.0%.

β grain size decreased from 1160 μm to 730 μm and β phase volume fraction increased from 22.7% to 35.0%. The α phase colony intercept length and α lath width were also refined from 3.2 μm to 2.2 μm and from 1.0 μm to 0.5 μm, respectively.2. Deposition using the GTA freeform fabrication system resulted in maximum grain refinement at an intermediate modulation frequency, with coarsening at higher frequencies. β grain size decreased from 1450 μm to 980 μm and β phase volume fraction increased from 14.7% to 24.6%. The α phase colony intercept length and α lath width was also refined from 2.9 μm to 2.2 μm and from 0.8 μm to 0.6 μm, respectively.3. Both modulated energy sources resulted in higher hardness values for Ti-6Al-4V alloy builds, with hardness increasing overall from 325 VHN to 356 VHN using modulation in the EBF3 machine and from 300 VHN to 321 VHN using modulation in the GTA machine.

The changes in hardness also resulted in enhanced isotropy in the builds, an indication of improved uniformity in552mechanical properties within the builds. A strong correlation was observed between hardness and α lath width using the Hall-Petch relationship.4. Dissimilarities between the two processes may be related to the differences in the power sources and wire consumables. The electron beam deposition method in the EBF3 system uses a rastering, high-density heat source and since the process occurs in high vacuum, requires the use of custom alloyed PCTWs for accurate deposition chemical composition. The CSM GTA system uses a traditional, relatively low-density GTA torch as the heat source and used conventional solid- core wire with unmodified chemistry. These variances may explain the different grain refinement-frequency trends and Hall-Petch parameters.

These variances may explain the different grain refinement-frequency trends and Hall-Petch parameters. AcknowledgementsThe authors acknowledge the support of the Advanced Materials and Processing Branch at NASA Langley Research Center and the use of the EBF3 system. One of the authors (S.M.) would also like to acknowledge the assistance of graduate students of the Colorado School of Mines Center for Welding, Joining and Coating Research Group.References[1][2][3][4][5][6][7][8]C. Hillier, "Powder-Cored Tubular Wire Development for Electron Beam Freeform Fabrication," Department of Metallurgical and Materials Engineering, Colorado School of Mines, Golden, 2010. B. Baufeld, O. V. d. Biest, and R. Gault, "Additive manufacturing of Ti-6Al-4V components by shaped metal deposition: Microstructure and mechanical properties," Materials & Design, vol. 31, pp. S106-S111, 2010. E. Brandl, V. Michailov, B. Viehweger, and C.

GRAPH GRAMMAR BASED APPROACH TO 3D PRINT AND ASSEMBLEFURNITURESulabh Gupta and Rahul RaiDesign Analytics Research and Technology (DART) Lab, Department of Mechanical and Aerospace Engineering, University at Buffalo-SUNY, Buffalo-NY, 14260AbstractA Plethora of user generated 3D models are available online. With rapid proliferation and diffusion of additive manufacturing machines in households, it has now become possible to download these virtual objects and print them out as physical parts. Although printing small size parts (within print volume of low cost 3D printers) is relatively an easy task, additive fabrication of large size parts (part volumes greater than print volume of low cost 3D printer) remains a challenging task for novice 3D printer users. In this paper the authors present a computational pipeline to 3D print large size 3D models that can be easily downloaded from online websites.

In this paper the authors present a computational pipeline to 3D print large size 3D models that can be easily downloaded from online websites. The pipeline essentially enables decomposition of large objects into smaller parts that can be 3D printed and then assembled. To assemble the printed parts a three-pronged approach is outlined. First, an interface based on graph grammar rules has been developed to generate assembly instructions. Second, an interactive segmentation of the desired 3D model is carried out using a Segmentation Guide Interface (SGI). SGI has been developed to assist a user to carry out component to sub-component segmentation. Third, we have also developed an interface that aids a user in printing small size pieces that can be printed in print volume of a commercial 3D printer (such as Makerbot®) and then assembled to create components that are too large to be printed in print volumes of low cost 3D printers.

We demonstrate the efficacy of developed pipeline by creating assembly instructions for multiple large sized 3D table models available online.IntroductionThe rapid development in 3D printing field has resulted in the beginning of a new era in personal fabrication. 3D printers are easy to use and are capable of printing almost any 3D CAD model. The widespread availability of 3D CAD models online has further stimulated personal fabrication. For example, online repositories have far more varieties of virtual furniture models than real ones in your nearby furniture store.Our goal is to enable individual users to 3D print their favorite 3D models by leveraging online shape repositories. In personal fabrication domain 3D printing until now has mostly been used for printing objects for trivial use or showpieces. In order to achieve large scale acceptance 3D printers should break free from their stereotypical personal fabrication use and print objects that have more practical utility.

In order to achieve large scale acceptance 3D printers should break free from their stereotypical personal fabrication use and print objects that have more practical utility. It has been observed that the maximum size of an object that a domestically affordable 3D printer can fabricate in one pass (the printing volume) is limited by practical considerations. This problem has curbed the large scale adoption of low cost 3D printers. Larger objects must therefore be printed as multiple separate parts and assembled.In this paper, we outline a computational pipeline to 3D print large size 3D models available online. Using the computational pipeline a user can semantically segment the 3D model into its component parts. Once the segmentation is completed the user is provided guidance in slicing the individual components into primitive shapes (sub components) which can be printed in the limited build volume of commercially available 3D printers.

The pipeline provides basic information to a user such as the number of sub-components required for each component, the shape of each sub-component and how to assemble the sub-components once they are printed. Additionally, we have developed a graph grammar based technique to generate assembly instructions. The generated assembly instruction joins components using off-the-shelf connectors found in furniture assembly domain.Our key contributions are:(1) We outline a computational pipeline that guides a user to 3D print large parts incommercially available 3D printers by segmenting them into smaller and printable sizes.(2) We have developed a set of new graph grammar rules that can be used for generatingassembly instructions to assemble furniture like tables.We demonstrate the validity of our computational pipeline by generating assembly instruction of 3D models of IKEA-style tables downloaded from Google 3DWarehouse, Princeton Shape Benchmark, and Polantis.

All our test models have arbitrary topology and structure. None of the 3D table models have the geometry or connector information needed for fabricating them.Related WorkIn the present section, three key areas related to the presented research namely (1) 3D printing (2) Graph Grammar, and (3) Mesh segmentation are briefly reviewed.3D PrintingWith the advancement of 3D printing we are witnessing the first stages of manufacturing democratization. It is believed that individual users will play a role in designing and creating their own products in the future [Gross 2007, Landay 2009]. Users can readily build their own customized products like plush toys [Mori, et.al 2007], chairs [Luo, et.al 2012], furniture [Lau, et al. 2011], garments [Umetani et al. 2011], Burr puzzles [Xin et al. 2011], and planar sections [McCrae et al. 2011, Hildebrand et al. 2012].

2011], garments [Umetani et al. 2011], Burr puzzles [Xin et al. 2011], and planar sections [McCrae et al. 2011, Hildebrand et al. 2012]. The current trend portends to revolutionize design and fabrication and could possibly give rise to a new class of creators and products [Mota, 2011]. We aim to propel this trend further. While the professional and higher end 3D machines are capable of producing large complex objects, the personal and less costly 3D printers are constrained by their print size and cost. The limited print volume of affordable 3D printers is the main cause of their limited usage. Efforts are been made to overcome these barriers.Luo et.al, [2012] has developed a new method called “chopper”. Chopper can be used to chop a given large objects into smaller pieces. These smaller pieces can be printed in 3D printersand assembled to create a large object. This method automatically generates male and female connectors on the cut surfaces.

These smaller pieces can be printed in 3D printersand assembled to create a large object. This method automatically generates male and female connectors on the cut surfaces. Once the object is printed, the user can assemble the parts without any assembly instructions. This method focuses on printing showpieces and the connectors generated do not provide much structural rigidity to the printed parts. In the computational pipeline outlined in this paper, a set of assembly instructions are generated to guide the user in assembling the 3D printed parts. The connectors used to assemble the parts are off-the-shelf connectors and the placement of these connectors provides structural rigidity to the final assembled part.Lau et.al, [2011] developed a graph grammar technique to create assembly instructions for virtual furniture models. In their work, the end user has the ability to use the assembly instructions to build furniture models with standard tools and wooden materials.

In their work, the end user has the ability to use the assembly instructions to build furniture models with standard tools and wooden materials. In the presented work, we guide the user to segment the 3D model into small parts. These segmented parts can be 3D printed in the limited build volume of low cost 3D printers. We then use graph grammar rules to generate assembly instructions for the printed models. The user can have little to no experience of working with tools.Graph GrammarGraph grammars have found applications in various fields such as concurrent systems, engineering, programming languages, and biology. It has been used to develop methods for understanding and modelling new architecture [Stiny 1980], cities [Parish, et.al 2001], buildings [Muller, et.al 2006] and details of facades [Wonka, et.al 2003]. Coffee maker grammar was one of first examples of using grammars for product design [Agarwal, et.al 1998].

Coffee maker grammar was one of first examples of using grammars for product design [Agarwal, et.al 1998]. Their grammar described a language that generates a large class of coffee makers. Shea et.al, [1997] presented a parametric shape grammar for the design of truss structures that uses recursive annealing techniques for topology optimization. Other engineering applications include lathe grammar [Brown, et.al 1997], grammar for machine design [Schmidt, et.al 1995], grammars for mechanical clocks [Starling, et.al 2003] and gear trains in [Starling, et.al 2005]. One of the interesting implementations of graph grammars is in the area of function-based design [Sridharan, et.al 2004]. In more recent applications graph grammar has been used to develop new design concepts to represent different topologies, configurations or shapes within a single search space [Rai, et.al 2011].

In more recent applications graph grammar has been used to develop new design concepts to represent different topologies, configurations or shapes within a single search space [Rai, et.al 2011]. Lau et.al, [2011] have used graph grammar techniques to generate assembly instructions for shelves, cabinets, and tables to build with standard wooden materials. We have developed graph grammar rules for generating assembly instructions for printing large size 3D printed objects. The novelty of our work is in the fact that a novice user can use our pipeline to easily assemble the 3D printed parts.Mesh SegmentationIn our work we segment the 3d model manually into its component parts. This is enabled through mesh segmentation techniques. Mesh segmentation techniques can be based on curvature and geodesic distances [Mangan, et.al 1999], dihedral angles [Shlafman, et.al 2002], Planarity and normal direction [Coheh-Stiner, et.al 2004, Attene, et al 2006], and Slippage [Gelfand, et.al 2004].

These attributes are sensitive to local surface features and to pose changes. Therefore, these techniques are not suitable for segmenting the same object in different poses. Topology-based approaches like spectral analysis [Liu, et.al 2004], average geodesic distance (AGD) [Hilaga, et.al 2001] and Reeb-graphs [Attene, et.al 2003] can be used for meshsegmentation of object in different poses. Nevertheless, they are vulnerable to topological and connectivity changes and do not distinguish well between shape differences.In our work we use Shape Diameter Function (SDF). The Shape-Diameter Function (SDF) is a scalar function defined on the mesh surface. It expresses a measure of the diameter of the object's volume in the neighborhood of each point on the surface.

The Shape-Diameter Function (SDF) is a scalar function defined on the mesh surface. It expresses a measure of the diameter of the object's volume in the neighborhood of each point on the surface. Using the SDF one can easily segment 3D models and perform part-retrieval [Shapira, et.al 2008].Basic Components of Computational PipelineFigure 1: Flowchart showing different steps of computational pipeline for printing and assembling parts bigger than the print volume of 3D printers.(Figure 1) illustrates the overall framework of the developed pipeline. The pipeline can accept 3D CAD models as input in any one of the following two ways: (1) A 3D CAD model created using CAD software like SolidWorks, CATIA, or PRO-E, or (2) A 3D CAD model obtained from online repositories such as Google Warehouse, Princeton Shape Benchmark, and Polantis.The Semantic Segmentation Interface (SSI) is used to segment the inputted 3D CAD models into component parts.

Each segmented part is saved individually. The SSI is based on SDF [Shapira, et.al 2008]. In order to carry out Semantic Segmentation using SDF vertex and face information is required. As a result it has to be ensured that the CAD model is in .obj format. If the input 3D model is not in .obj format then MeshLab software can be used to convert the 3D CAD file into the required (.obj) format.GraphSynth (a graph grammar software) is used to create assembly instructions to assemble the segmented parts. For assembly instruction generation we have developed twenty- four rule graph grammar that defines different ways to assemble tables (Table 1). By simply executing different combinations of grammar rules, assembly instructions for a variety of tables can be generated.

By simply executing different combinations of grammar rules, assembly instructions for a variety of tables can be generated. A user skilled in wood working can use the instructions generated by GraphSynth to make the object using wood and simple fasteners.In our proposed approach it has been assumed that a user has little or no experience with wood working tools. As a result our computational pipeline enables a not-so-skilled user to 3D print the parts. 3D printers require the input files to be in .stl format as a result the segmented components that are in .obj format have to be converted to .stl format. .obj to .stl format conversion can also be carried out using MeshLab. In case segmented 3D components do not fit in the build volume of the 3D printer they are further segmented. We further segment the components into sub-components using slicing software called netfabb Basic.The printed sub-components are assembled to create components.

We further segment the components into sub-components using slicing software called netfabb Basic.The printed sub-components are assembled to create components. The components are then assembled to create the overall object by using the assembly instructions generated by graph grammar interface. The tool components and the associated methods are described in detail in the following sub-sections.Semantic Segmentation Interface (SSI)(a)(b)(c)(d)Figure 2: (a) Semantic Segmentation Interface, (b) Semantic segmentation carried out by drawing a boundary around the part to be segmented, (c) Segmented table top, (d) The segmented table into its constituent partsThere is lack of assembly information in the 3D models available online and most of the 3D models are stored as a single part. The Semantic Segmentation Interface (SSI) is used to semantically segment 3D CAD models into component parts (Figure 2a).

The Semantic Segmentation Interface (SSI) is used to semantically segment 3D CAD models into component parts (Figure 2a). SSI is based on a volume-based shape-function called the shape-diameter function (SDF). The SDF is a scalar function deﬁned on the mesh surface. In essence it represents a measure of the diameter of the object’s volume in the neighborhood of each point on the surface [Shapira et al 2008]. In order to carry out Semantic Segmentation using SDF, vertex and face information is required. As a result it has to be ensured that the input 3D CAD model is in .obj format. If the inputted 3D CAD model is not in .obj format then MeshLab can be used to convert it into .obj format. User through interaction with SSI segments the 3D CAD model by drawing a box around the part (Figure 2b). The user can also select different colors to color segmented parts in different colors. This helps in differentiating different segmented parts.

The user can also select different colors to color segmented parts in different colors. This helps in differentiating different segmented parts. Each part can then be exported separately.Graph Grammar for generating assembly instructionsGraph grammars are comprised of rules for manipulating nodes and arcs within a graph. In our computational pipeline the graph grammar rules specify a formal language for generating assembly instructions from an initial user defined graph (seed graph). The development of these rules encapsulates a set of valid operations that can be used to assemble different parts of a table. Through the application of each grammar rule the current state of graph is transformed into a new state, incrementally evolving towards a desired solution.A typical graph grammar rule is comprised of a left-hand side (LHS) and a right-hand side (RHS). The LHS contains the conditions, upon which the applicability of a rule is determined.

The LHS contains the conditions, upon which the applicability of a rule is determined. Accordingly, the LHS describes the state of the graph for a particular rule to be applicable. The RHS, on the other hand, contains the resulting graph transformation. It describes the new state of the graph after the application of the rule. We have developed a twenty four-rule graph grammar that defines ways to assemble different parts of a table (Table 1).

The generation of assembly instruction starts with the user defining the initial contact points of the table. This is a simple step and the user inputs contact point information in the form of a seed graph (Figure 3). Once the seed graph is defined, in the following steps different graph grammar rules are used to automatically recognize differentconnections and appropriate connectors. This continues until all the connectors for the whole assembly are identified.A complete sequence of application of different assembly instruction graph grammar rules to create assembly instructions for a table shown in (Figure 3) has been illustrated in (Figure 4). In order to generate assembly instructions graph grammar approach starts with a seed graph which is defined by the user. This results in the seed graph to become LHS in step 1. After this stage the process of recognize, choose, and apply is invoked in an iterative manner resulting in a new LHS and RHS at each step.

This results in the seed graph to become LHS in step 1. After this stage the process of recognize, choose, and apply is invoked in an iterative manner resulting in a new LHS and RHS at each step. The assembly instruction for the table example shown here can be derived by application of rule sequence {1, 1, 1, 1, 6, 6, 6, 6}. Since tables have repetitive parts like legs we notice that the same rules have been applied more than once. The graph grammar interface can generate assembly instructions for a wide range of tables.Table	
 TopApplicable rule 1, 6C1LegsApply rule 1 four timesC2Applicable Rule 6ShelfApply rule 6 four timesTERMINATEFigure 4: Application of rule sequence {1, 1, 1, 1, 6, 6, 6, 6} creates the assembly instructions for the example table from initial seed graphSegmentation Guide Interface (SGI)A single component can be too big to 3D print at one go.

The idea behind the development of SGI is to guide the user to carry out further segmentation of the segmented components. SGI enables the user to compare the build volume of the 3D printer to the volume of the component. If the print size of the component is smaller than the print volume of 3D printer then the part is ready to be printed. Otherwise SGI provides information regarding the shape and number of sub-components that the component should be segmented into. SSI also provides the user with visual information regarding type of connectors and how to use those connectors to join the sub-components.The Segmentation Guide Interface has been developed in Visual C#. Visual Studio supports Visual C# with a full-featured code editor, compiler, project templates, designers, code wizards, a powerful and easy-to-use debugger, and other tools.

Visual Studio supports Visual C# with a full-featured code editor, compiler, project templates, designers, code wizards, a powerful and easy-to-use debugger, and other tools. The .NET Framework class library provides access to many operating system services and other useful, well-designed classes that speed up the development cycle significantly.A sequence illustrating the interaction with SGI is presented below:1) Prompt the user to enter the build dimensions printer. of Maximum printable area and volume is calculated (Figure 5a). 2) The user is then prompted to enter the shape of the table top. A choice of three shapes is displayed (Figure 5b).the(a)(b)3) Once the user selects the shape of the table top. The user is prompted to enter the dimensions of this information area and the volume of the table top are calculated (Figure 5c). the(c)4) A list of possible connector types is displayed (Figure 5d) to join the sub-components.

the(c)4) A list of possible connector types is displayed (Figure 5d) to join the sub-components. A (.gif file) appears to help the user visualize the selected connector. button.5) The user then clicks on the The segmentation interface responds by displaying the number and shape of each sub-component the user needs to generate for creating the overall component (Figure 5e). that(d)(e)Figure 5: Segmentation Guide Interface working for segmenting the Table top6) Similar steps are followed to segment the Legs of the table.the segmentation guide interface are passed on to netfabb Basic software. Netfabb Basic is very simple to use software and is often used to segment parts that are bigger than the build volume of 3D printers. netfabb Basic is a freeware for handling .stl files. The freeware includes advanced model fixing, measurement and quality management.

netfabb Basic is a freeware for handling .stl files. The freeware includes advanced model fixing, measurement and quality management. The freeware also includes a basic slicing module and assists the user in data preparation and 3D Printing.Thesuggestionsprovidedbybrowsing,STLFigure 6: Semantically segmented components being segmented into sub-components using netfabb Basic.The .stl files of the components are imported into netfabb Basic and cut option is used to slice portions of the model (Figure 6). The number of pieces and the shape to be segmented is based on feedback from SGI. Each single piece is exported separately. User can often save time if the parts repeat themselves; in such cases only one sub-component is exported. These parts are then separately exported to the 3D printer’s software to begin printing.ResultsThe outlined computational pipeline is tested on a set of five table CAD models.

These parts are then separately exported to the 3D printer’s software to begin printing.ResultsThe outlined computational pipeline is tested on a set of five table CAD models. The CAD models for these tables were downloaded from Google 3D Warehouse, Polantis, and Princeton Shape Benchmark websites. An illustrative example shows how the computational pipeline could be used to 3D print and assemble a simple table (Figure 7).	
 The run time of the entire proposed approach depends on the user familiarity with the different software involved.

The software involved is very easy to learn and use.(a)(b)(c)(f)(e)(d)(g)Figure 7: (a) Downloaded model is segmented into its constituent components using SSI, (b) SGI is used to determine the shape and number of sub-components to be created, (c) GraphSynth is used to generate assembly instructions, (d) Segmented sub-components are 3D printed, (e) 3D printed sub-components are assembled, (f) The final table is assembled using assembly instructions generated by GraphSynth. and (g) the final assembled table(b)(c)Assembly instructions generated by the piepline for four other tables is shown in Figure 8.(a)(b)1397(c)(d)Figure 8: Assembly Instruction for IKEA style furnitureConclusionWe have developed a computational pipeline to enable a user to 3D print parts that are larger than the print volume of 3D printers.

To demonstrate the utility of the developed pipeline, we 3D print segmented parts of five different tables using MakerBot®	
 Replicator™ 2X and assemble these parts using the assembly instructions generated by the computational pipeline. Some printed parts had warping problems resulting in significant distortion of the printed parts. Sometimes post processing of parts like enlarging hole size and hammering certain joints is required. The finish quality of the parts is dependent on the size and alignment of the print. For example, the segmented legs when printed in upright position gave better fit compared tohorizontal prints. We believe that with further advancement in 3D printing processes these issues can be easily addressed.To make our framework more general, one can extend the presented pipeline to work with any arbitrary 3D model. This could be achieved by extending the initial set of 24 rules presented in this paper.

This could be achieved by extending the initial set of 24 rules presented in this paper. The graph grammar rules defined in the pipeline have been derived manually. Automatic graph grammar rule generation for an arbitrary 3D model is an avenue for future research. The interfaces used in our approach require user interaction to carry out manual segmentation. It would be beneficial to enable automatic segmentation of the components and the sub-components.References[1] Gross, M. "Now more than ever: computational thinking and a science of design." Japan Society for the Science of Design 16, no. 2 (2007): 50-54.[2] Landay, James A. "Technical perspective Design tools for the rest of us."Communications of the ACM 52, no. 12 (2009): 80-80.[3] Mori, Yuki, and Takeo Igarashi. "Plushie: an interactive design system for plush toys." In ACM Transactions on Graphics (TOG), vol. 26, no. 3, p. 45. ACM, 2007.[4] Luo, Linjie, Ilya Baran, Szymon Rusinkiewicz, and Wojciech Matusik.

 ME(cid:6)TOR: A High School Education Program in Systems Engineering and Additive ManufacturingDavid W. Rosen1, Dirk Schaefer1, Daniel Schrage21School of Mechanical Engineering, Georgia Institute of Technology 2School of Aerospace Engineering, Georgia Institute of TechnologyAbstractThe Defense Advanced Research Projects Agency (DARPA) is sponsoring the MENTOR program as the outreach part of its Adaptive Vehicle Make suite of programs. Georgia Tech has been awarded a contract to involve up to 1000 high schools after 4 years in a series of prize challenges. A web-based collaborative design-manufacturing infrastructure will be developed that integrates CAD, CAE, design-for-manufacturing, and CAM software tools with a network of 3D printers and other manufacturing resources. In distributed teams, students will design, fabricate, and construct electro-mechanical systems (e.g., ground vehicle robots) to perform complex tasks.

In distributed teams, students will design, fabricate, and construct electro-mechanical systems (e.g., ground vehicle robots) to perform complex tasks. Many parts they design will be fabricated on 3D printers that are located in high schools or nearby sites. A project objective is to have students learn about collaborative design, advanced manufacturing, and new product development practices and to become excited about pursuing technology-based careers.1 DARPA Manufacturing Experimentation and Outreach (MENTOR)The DARPA Manufacturing Experimentation and Outreach (MENTOR) BAA was issued October 18, 2010 by the Tactical Technology Office (TTO). The total amount of money to be awarded was $10.0 million with multiple awards anticipated. The MENTOR effort is part of the Adaptive Vehicle Make (AVM) program portfolio and “…is aimed at engaging high school students in a series of collaborative distributed manufacturing and design experiments.

The overarching objective of MENTOR was to develop and motivate a next generation cadre of system designers and manufacturing innovators, and to ensure that high school-age youths are exposed to the principles of modern prize-based design and foundry-style digital manufacturing” (quotes are taken from the BAA).The end vision for MENTOR was the “…development of user-friendly, open-source tools to enable the utilization of conventional social network media (e.g., Facebook apps) for the purpose of collaborative distributed design and manufacturing across hundreds of sites and thousands of users.” This capability will be accompanied by the “deployment of an inexpensive, heterogeneous set of digitally-programmable manufacturing equipment (e.g., 3D printers for various materials) to 1,000 high schools globally.” Prize-based design and manufacturing challenges would then “enable clusters of schools to team and compete against one another in the development of cyber-electro-mechanical systems of moderate complexity such as go carts, mobile robots, small unmanned aircraft, etc.”Three teams were awarded contracts for the MENTOR program in Fall 2011, including one from Georgia Tech.

Each is a 1-year contract, with options for succeeding years. The Georgia Tech team is hopeful of being renewed to continue working on the project.2 The Georgia Tech ApproachGeorgia Tech being the U.S. largest producer of engineers has the resources and mechanisms in place to meet the vision and satisfy the objectives that DARPA seeks in its MENTOR Program. Faculty from three schools in the College of Engineering: Aerospace Engineering (AE), Electrical & Computer Engineering (ECE), and Mechanical Engineering (ME), who have had extensive design & manufacturing and outreach interest and experience, teamed together, along with other key Georgia Tech personnel to provide the necessary leadership. An illustration of the GT MENTOR organizational structure and element leads is provided in Figure 1.Figure 1. GT MENTOR Organizational Structural Elements.Dr. Dan Schrage, Professor and Director of the Integrated Product Lifecycle Engineering (IPLE Laboratory in the School of AE, and Dr.

GT MENTOR Organizational Structural Elements.Dr. Dan Schrage, Professor and Director of the Integrated Product Lifecycle Engineering (IPLE Laboratory in the School of AE, and Dr. David Rosen, Professor and Director of the Rapid Prototyping and Manufacturing Institute (RPMI) in the School of ME, serve as the Co- Directors /Co-Principal Investigators (Co-PIs) for the GT MENTOR Program. Dr. Rosen, along with Professor Dirk Schaefer, School of ME, lead the Integrated Design & Manufacturing Infrastructure (IDMI) element. Mr. Dale Atkins, Senior Research Engineer, and Mr. Srujal Patel, Research Engineer, in the School of AE lead the MENTOR Academy Central (MAC) element which overseas the day-to-day activities of MENTOR, as well as planning and interfacing its self-sustainment. Dr. Schrage, along with Dr. Bonnie Ferri, Professor and Associate Chair of the School of ECE, and Dr.

Dr. Schrage, along with Dr. Bonnie Ferri, Professor and Associate Chair of the School of ECE, and Dr. Tom Collins, a Senior Research in the Georgia Tech Research Institute (GTRI), along with a joint appointment in the School of AE, conduct and oversee the Prize Challenge Development (PCD) and the Prize Challenge Activities (PCA) elements. Joining them in this capacity is Mr. Olivier Ammoun, a Senior Research Manager with DassaultSystemes. Dassault Systemes (DS) provided to GT MENTOR several of the capabilities that the DARPA MENTOR Program sought. First, they are the world’s leader in introducing modern Product Lifecycle Management (PLM) knowledge for outreach education activities. The DS PLM Discover middleware architecture was specifically developed for education outreach activities to illustrate and implement how CAD, CAE, and CAM can be applied for cyber- electro-mechanical systems of moderate complexity, as specifically identified in the DARPA MENTOR BAA.

The DS PLM Discover targets high school and undergraduate non-expert curricula as illustrated in Figure 2. Being web-based it strongly supports the collaborative national and international environment sought by DARPA to engage 1,000 high schools worldwide within four years.Figure 2. Specific Educational Targets for DS PLM Discover.The fact that DS already supports in France an educational prize-based competition with more than 1,000 high schools participants from around the world, The Formula One Race Car Competition, greatly lowers the risk for the GT MENTOR Team to reach the DARPA target of 1,000 high schools for its prized-based challenges by the end of four years. CATIA V5 is the basis CAD model used in this competition.Another key DARPA MENTOR desired capability that DS provides the GT MENTOR Team is user-friendly, open-source tools to enable the utilization of conventional social network media.

DS has provided the GT MENTOR Team their See what You mean (SwYm) social network tool, which has already been evaluated in the Georgia Tech 2011 Engineering Summer Camp and will be extensively evaluated in the 2012 Engineering Summer Camp in July 2012. Some sample SwYm results from the 2011 Summer Camp, prior to MENTOR initiation, are provided in Figure 3. It is felt that this evaluation of social networking is innovative and unique. It may also be the first time that social networking has been used and evaluated in a prize- challenge competition simultaneously executed by collaborative team members from two different camps at two different locations.Figure 3. Sample SwYm Results from 2011 Summer Camp.The Evaluation & Assessment (EA) task in Figure 1 is led by Dr. Meltem Alemdar, Assistant Director for Evaluation in the Georgia Tech Center for Education Integrating Science, Mathematics, and Computing (CEISMC). Dr.

Meltem Alemdar, Assistant Director for Evaluation in the Georgia Tech Center for Education Integrating Science, Mathematics, and Computing (CEISMC). Dr. Alemdar’s work to date has been preparing the GT MENTOR Evaluation Plan. However, the bulk of her time has been spent preparing the Institute Review Board (IRB) Proposal for the GT MENTOR Program, which is still undergoing review.Another key element of the GT MENTOR Organizational Structure in Figure 1 is the High School Clusters (HSC). The HSC element is lead by Dr. Kelly Griendling, GT MENTOR Outreach Coordinator. Her responsibilities have been informing high schools about GT MENTOR and identifying the approximately ten development high schools for the first two years of MENTOR development. Dr. Griendling has a large, strong group of GT MENTOR development high schools for the first year, as illustrated in Figure 4. As can been seen 15 development high schools have been identified with five more having expressed interest.

As can been seen 15 development high schools have been identified with five more having expressed interest. With a commitment to exercise the second year of the DARPA MENTOR development phase, it is felt that 40-50 high school clusters could be provided worldwide by the end of the third year. With 8 to 10 high schools per cluster, it is felt that approximately 400 high schools could be involved in the GT MENTOR Program by the end of the third year, greatly exceeding the 100 high schools specified in the DARPA MENTOR BAA. With funding for the year three option, it is strongly felt that the 1000 high schools required by end of the fourth year could easily be obtained. This estimation is strongly supported by the GT MENTOR partner, Dassault Systems, who has successfully supported the Formula One Race Car Competition in France. Also, they have gauged strong interest in the GT MENTOR Program from the numerous high schools they are in contact with.Figure 4.

Also, they have gauged strong interest in the GT MENTOR Program from the numerous high schools they are in contact with.Figure 4. GT MENTOR Development High School Clusters.The final element identified in the GT MENTOR Organizational Element Chart, Figure 1, is Prize Based Activities (PCA). As identified, Dr. Tom Collins and Dr. Bonnie Ferri have led two Prized Challenge development efforts supported by the MAC graduate and undergraduate students from by the IPLE Laboratory. Dr. Collins led the Wind Turbine Prize Challenge, while Dr. Ferri led the Ground Robot Prize Challenge. Initial Prize Challenge concepts were developed from baseline LEGO NXT kit as illustrated in Figure 5. The Wind Turbine and Ground Robot were evaluated in the 2011 Engineering Summer Camp. While they use the LEGO NXT kits as baselines, the intent is for the high school teams to Co-Create, Design, Build and Operate (CDBO) non-standard parts.

While they use the LEGO NXT kits as baselines, the intent is for the high school teams to Co-Create, Design, Build and Operate (CDBO) non-standard parts. For the Wind Turbine this consists of new turbine blades to increase the efficiency for power generation as illustrated on the bottom of Figure 6. For the Ground Robot the CDBO evolution from the Virtual to the Physical for an external non-standard cover is illustrated at the top of Figure 6. These initial Prize Challenges were developed and evaluated during the 2011 Engineering Summer Camp. The 3D pprinter utilized was from Stratasys. Both Stratasys and 3D Systems are GT MENTOR partners and represent two of the largest 3D printer companies.

The 3D pprinter utilized was from Stratasys. Both Stratasys and 3D Systems are GT MENTOR partners and represent two of the largest 3D printer companies. Thus, GT MENTOR has responded to another of the DARPA MENTOR BAA Objectives: deployment of an inexpensive, heterogeneous set of digitally-programmable manufacturing equipment (e.g., 3D printers for various materials) to 1,000 high schools globally.The technical approach for the GT MENTOR Program is illustrated in Figure 7. It provides a CDBO framework for collaborative prize challenge development and implementation. It includes the four major life cycle functions for design, development and testing of cyber-electro- mechanical systems. The arrows indicate how the PLM Discover architecture can be used to exchange, store, distribute and re-use the vital product and process data between distributed highschool teams for decision-making within their clusters.

It also provides the high school students and teachers an understanding of the need for integrated design and manufacturing tradeoffs. In addition, it provides in an easily, understandable way the need for cycle time reduction, a key objective for the DARPA AVM Program, for which MENTOR is supposed to be a microcosm.Figure 5. Initial Prize Challenge Development.Figure 6. Wind Turbine and Ground Robot Prize Challenges.Figure 7. GT MENTOR Technical Approach.Each summer Dr. Schrage offers a 2 week summer camp on engineering systems design to high school students. Students learn CATIA, collaboration, systems engineering, and 3D printing. Students work in teams of 2-4. During the 2011 Engineering Summer Camp, teams were composed of 2 students at Georgia Tech and 2 more at the University of Detroit-Mercy. Students worked on either the wind turbine or ground vehicle projects.

Students worked on either the wind turbine or ground vehicle projects. The technical approach illustrated in Figure 7 was initially evaluated for the 2011 summer camp and proved to be quite successful. The 2012 Engineering Summer Camp will be July 9-20, 2012 and will include approximately forty high students and ten high school teachers. Many of the students and teachers will be from the GT MENTOR development high schools. In addition, students from six states and one foreign country have also enrolled for the 2012 Summer Camp, many coming with the desire to host their own summer camps in their regions in 2013.One of the key MENTOR challenges that the GT MENTOR Team has addressed is sustainment. A necessary ingredient for sustainment to occur is to engage high school teachers in professional education and participation in the summer camps, which will help them become the true mentors. Another is to provide a mechanism that will help them upgrade their high school curriculum.

Another is to provide a mechanism that will help them upgrade their high school curriculum. With this in mind, the GT MENTOR Program is partnering with the Southern Region Education Board (SREB), particularly with their Preparation For Tomorrow (PFT) initiative. The SREB activities are focused in 20 states, mostly in the Southeast, but spanning from Texas to Delaware. The PFT initiative is to introduce project-based learning through state focused projects over the next five years. Twelve of the states have agreed to incorporate four project-based learning courses (one per year) into their high school curricula, including Georgia, which selected advanced manufacturing as their focus.The GT MENTOR Program is working with the SREB PFT to include prize-based challenges into the four project-based learning courses. This has caused the GT MENTOR Program to reconsider its prize challenge progressions from Grades 9-12 grades.

This has caused the GT MENTOR Program to reconsider its prize challenge progressions from Grades 9-12 grades. For Grade 9,the CDBO approach for new turbine blades, similar to Figures 5 & 6, would be followed. It is felt that use of the LEGO NXT kits is a good start for most prize challenges and project-based learning topics. However, it is felt that the progression should end up with something that is visible and useful to the school. Thus, the objective for the Wind Turbine is to end up with a small workable wind turbine could be mounted to the school or a pole of some type and demonstrated power efficiency enhancement. Some restaurants and small businesses already use this approach. Thus, in Grade 10 a small wind turbine would be selected and a CDBO approach to mounting it would be the project.

Some restaurants and small businesses already use this approach. Thus, in Grade 10 a small wind turbine would be selected and a CDBO approach to mounting it would be the project. In Grades 11 & 12, the small wind turbine would be installed, operated and tested to determine what energy enhancements can be made.Similar state focused projects will be identified for ground and aerial robots. A ground robot that could assist the high school janitors in doing their jobs could be another application. An aerial robot that could monitor school security from outside could be another application.3Integrated Design-Manufacturing Infrastructure (IDMI)An integrated design and manufacturing infrastructure has been designed and implemented to fulfill requirements for students to gain access to the collaboration environment and design and manufacturing tools.

In this subsection, we provide an overview of the system and describe a variety of its key feature sets that have been implemented.The innovative approach for this work was to provide an IDMI that supports smooth transitions between design and manufacturing resources in order to accommodate timely design and manufacturing tradeoffs and associated processes. The goal was to enable these design and manufacturing transitions in a highly open and collaborative environment.

The goal was to enable these design and manufacturing transitions in a highly open and collaborative environment. As a result, the system utilized open-source tools and technologies for the vast majority of its feature sets and focused on tools that enable the seamless integration of collaboration technologies [1].While investigating the various software tools required for the IDMI implementation, the following criteria were established to guide our evaluations: 1) the code must be open source so that we can modify existing features and/or add new features when needed; 2) the systems must be built on top of ‘standard’ and well-established technologies; for example, we consider systems based on Java programming technologies, PHP, etc to be well-established.

3) the systems must be flexible in terms of deployment and user requirements, utilize common operating systems, and support standard HTTP servers, web browsers, database systems (e.g., MySQL).The Moodle learning management system was selected to function as the underlying centralized interfacing portal of the IDMI. Moodle is a free web-based application used by educational entities around the world to create online learning systems. Moodle is based on standard technologies such as PHP, HTML, CSS, and AJAX. Moodle can be implemented on top of a traditional LAMP stack, which includes a Linux operating system, Apache web server, MySQL database, and PHP. These LAMP components along with Moodle are open source.The IDMI has a centralized interfacing server (CIS) that hosts the primary IDMI portal and other IDMI resources such as collaboration tools, middleware services, and cloud file systems, which - taken together - form a small cloud-based design-manufacturing environment [2,3].

Figure 8 illustrates the main landing page for the IDMI portal.The landing page has a 3-panel design along with a header and footer. The left and right panels are based on a ‘block’ layout design whereby different functionality can be added as block elements that contain features such as site navigation, news, RSS feeds, and calendars. Themiddle pane serves to illustrate the four main categories of resources provided by the IDMI: Design & Manufacturing resources, Collaboration resources, Learning & Education resources, and Teaching resources. Users gain access to the system by clicking the ‘Login’ hyperlink located at the top right-hand side of the landing page. Upon successful login, the user will be directed to his/her home page.Figure 8: Main IDMI portal landing page.The system uses role-based access control. The content displayed to the user will depend on the user’s role. However, various types of content are displayed independently of role.

The content displayed to the user will depend on the user’s role. However, various types of content are displayed independently of role. For example, site administrators will have access to the features provided by the site administration block. Non-administrative users will not see this content. The generic content that all users see, however, are the links to the IDMI’s resources defined by the system’s portal information architecture along with the file management block, user list block, and site navigation block.Site administrators access tools for system-wide management via the site administration block. These tools are used to manage system users, IDMI resources, student grades, location settings, language settings, new plugin features, security settings, site appearance, front-page settings, server settings, system event reports, and portal development.

After adding a new user to the system, the site administrator assigns one or more roles to the user.The IDMI portal has been designed to support four types of roles, including: Student, Assistant, Teacher, and Manager. These roles determine read, write, and modify permissions for the content available from within the portal. These roles also determine contextual views available within the site’s pages. For example, only managers have access to site administrator tools and these tools are not viewable for non-managers.In terms of web systems, an information architecture defines the high level categorization and organization of information, resources, and services available to users of a web-based system. For the IDMI, the information architecture is organized around four main categories: Design & Manufacturing resources, Collaboration resources, Learning & Education resources, and Teaching resources.

For the IDMI, the information architecture is organized around four main categories: Design & Manufacturing resources, Collaboration resources, Learning & Education resources, and Teaching resources. The elements and functionalities of these four categories will be described by the following.Figure 9 illustrates the IDMI Design & Manufacturing interface. From the design and manufacturing resources interface, users have access to CATIA V6 CAD software provided via a CITRIX web portal interface. The CATIA V6 interface is illustrated by Figures 10 and 11 respectively. From this interface, users also have access to the various additive manufacturing tools that have been incorporated into the IDMI, such as AM-Select and AM-Request, which will be discussed in the next section.

Other features available from the Design & Manufacturing interface include collaboration tools such as RSS feed blocks, news and event broadcast blocks, system tools such as log, activity, and participation reports, and generic tools such as forum topic search functionality. Lastly, each top-level category interface has its own dedicated news forum whereby users can collaborate in real time and/or asynchronously.Figure 9: IDMI Design & Manufacturing interface.Figure 11 illustrates the Collaboration interface of the IDMI. The following technologies have been integrated into the IDMI for both real time and asynchronous collaboration: Chat rooms, Forums, Wikis, Workshops, Blogs, and Agendas. The blogs, forums, and wikis that have been implemented in the IDMI are state of the art and include the capability via web-based WYSIWYG editors to input text and images, which provides for better collaboration during online design processes.

Figure 12 illustrates the blogging functionality that has been implemented for the IDMI.A number of external collaboration systems have been integrated into the IDMI. In particular, the IDMI interfaces with the Wiggio video telecollaboration (VTC) and virtual meeting tool, and a collaborative cloud file system based on Dropbox has been implemented.Figure 13 illustrates the video VTC capabilities of the Wiggio tool. In this scenario, two engineers are collaborating via VTC technology. They both have access to shared files via the cloud file repository of the IDMI and its dropbox functionality. In this scenario, the team is collaborating on a particular design. One team member located in Atlanta, GA has added his STL file to the repository.

In this scenario, the team is collaborating on a particular design. One team member located in Atlanta, GA has added his STL file to the repository. The other team member located in Savannah, GA is retrieving the file from the cloud repository and is preparing to have the file processed by a 3D printer.Figure 10: CATIA V6 CAD software accessed by users via the IDMI portal.Figure 11: IDMI interface for the primary collaboration tools.The learning resources interface of the IDMI provides access to students for components such as training materials, course curriculum, and wiki materials. The teaching resources interface of the IDMI provides resources for teachers who use the system.

The teaching resources interface of the IDMI provides resources for teachers who use the system. Other thancollaboration forums for teachers using the IDMI, the teaching interface also includes tools such as plagiarism detection, course management, grade management, scheduling, agendas, and outcome reporting, just to name a few.Figure 12: An IDMI Blogging session.An integrated design and manufacturing infrastructure has been designed and implemented. The innovate approach for this work was to provide an IDMI that supports smooth transitions between design and manufacturing resources in order to accommodate timely design and manufacturing tradeoffs and associated processes. The goal was to enable these design and manufacturing transitions in a highly open and collaborative environment.

The goal was to enable these design and manufacturing transitions in a highly open and collaborative environment. As a result, the system utilized open-source tools and technologies for the vast majority of its feature sets and focused on tools that enable the seamless integration of collaboration technologies. An overview of thesystem has been given in this section along with a description of a variety of its key feature sets that have been implemented.Figure 13: Wiggio integration and functionality within the IDMI.4 Additive Manufacturing ToolsA set of design-manufacturing software tools were developed to enable students to learn about AM, how to design parts to be fabricated using AM, and how to select appropriate AM processes, machines, and materials, and how to request parts to be fabricated. Alpha versions of the tools, called AM-Select, AM-Request, and AM-Manufacturable, have been integrated into the IDMI framework, as shown in Figure 9, and will be described here.

Alpha versions of the tools, called AM-Select, AM-Request, and AM-Manufacturable, have been integrated into the IDMI framework, as shown in Figure 9, and will be described here. Each application was developed in the C++ language and is web-enabled.The proposed workflow through the design-to-manufacture tools is shown in Figure 14. Starting in a CAD system, such as CATIA, students create their part designs and save them as STL files. From there, they can learn about the manufacturability characteristics of their designs using the AM-Manufacturable application. This will highlight difficult-to-fabricate features and provide feedback about part size, feature orientation, part orientation in the machine, and other related aspects. Students then have the option of going back to the CAD system to redesign their parts.

Students then have the option of going back to the CAD system to redesign their parts. The AM-Select application reads the part STL file and makes recommendations of processes, machines, and materials, which are passed to the AM-Request application to make the actual request for fabricating the part. Clicking on one of the boxes in Figure 14 causes the software application to launch.Figure 14: Workflow from CATIA to printed parts.4.1 AM-ManufacturableThe application AM-Manufacturable reads a part STL file and provides manufacturability feedback to the student about their design. A mock-up of part of the proposed user interface is shown in Figure 15. The interface gets the student thinking about how their part will be fabricated in a 3D printer by asking them about the part’s orientation in the machine.

The interface gets the student thinking about how their part will be fabricated in a 3D printer by asking them about the part’s orientation in the machine. By changing the build orientation, AM-Manufacturable will update the “Build Height” and “Support Structure Volume” entries in the interface.Feedback will be provided on features that are too small to build. Optionally, the user can select a 3D printing process and the database will provide the corresponding smallest feature capability. Or, the user can just enter the feature size with which to evaluate the part. A novel feature size computation algorithm is executed to identify faces that are close to one another. From there, the feature is recognized and its size is determined. Small features are highlighted in the STL part representation and displayed in the CAD system, an example of which is shown in Figure 16, where small features are highlighted in red.Additional functionality is planned for part size evaluation.

If a part is too large to fit into a 3D printer, the user will be notified. Suggestions for how to divide the part into an assembly will be provided.Figure 15: Mock-up of the first AM-Manufacturable interface screen.4.2 AM-SelectUsing the AM-Select application, a user can identify and compare candidate manufacturing processes based on build time and cost estimates that are computed from approximate geometric information about the part. Within each class of manufacturing processes, the user can also identify specific machine types and materials that meet his/her needs. AM-Select can be executed on approximate geometric and size information, specifically the part’s bounding box and its estimated volume, which enables process selection at early design stages.

AM-Select can be executed on approximate geometric and size information, specifically the part’s bounding box and its estimated volume, which enables process selection at early design stages. Alternatively, AM-Select can be run using the part’s STL file so that more precise estimates of build time and cost can be computed.Thin rib and cut section Thin rib and cut sectionSmall holes Small holesFigure 16: Identified small features in two example parts. Small features are indicated by the red surfaces.The build time model is based on a generalized parameterization of AM processes that applies to laser-based scanning (Stereolithography, powder bed fusion), filament extrusion (fused-deposition modeling), ink-jet printing, and mask-projection processes.

Build time estimates were tested by comparing them to the measured build time of parts in fabricated using Stereolithography, ink-jet printing, and fused deposition modeling processes.The overall cost is estimated by accumulation of four sub-costs: machine purchase cost (P), machine operation cost (O), material cost (M), and labor cost (L) as shown in Eq.1. Each term is given by a corresponding model. The machine operating cost model involves the estimation of part build time; separate models are used for laser scanning, fused deposition modeling, and ink- jet printing processes, but all share a common structure. More details can be found in reference [4].(1)A database of processes, machines, and materials was developed for the 3D printers currently commercially available from 3D Systems, Stratasys, and Objet. The database format is straight-forward enabling easy updating when new machines and materials become available.

The database format is straight-forward enabling easy updating when new machines and materials become available. The database structure was developed as part of the DARPA iFAB project, ensuring its comprehensiveness and robustness.The AM-Select application has two primary screens to support user interaction. In the first, the user provides either the STL file name or the part’s bounding box and estimated volume. Then, the user specifies design specifications on surface finish, geometric accuracy, strength, and stiffness. In all cases, only qualitative values (e.g., low, medium, high) are needed. The first screen is shown in Figure 17.Figure 17: Information input screen for AM-Select.The second screen shows the results of the build time and cost estimates. Of the machine time and material combinations listed, the user can select one, some, or all listed.

Of the machine time and material combinations listed, the user can select one, some, or all listed. When the user hits the “Submit Build?” button, the selected machines and materials are written to a project file in the IDMI that downstream applications can access. An example screen is shown in Figure 18.4.3 AM-RequestThe application AM-Request has a simple purpose: enabling the user to select one machine and material to build his/her part. This selection is for machines that are currently available to the user to use. Upon selection, an email message is sent to the machine operator requesting the part be built. The part STL file is attached to the email. The input to AM-Request is the file that was saved by the AM-Select application, enabling a smooth work-flow from one application to another.The AM-Request user interface is shown in Figure 19. When the “Retrieve Next Order” button is hit, the AM-Select file is read.

When the “Retrieve Next Order” button is hit, the AM-Select file is read. A database of available machines, materials, and machine locations is read and compared with the selected machines and materials from AM- Select. At present, this database of available machines was created manually, but in the future, this database will be automatically generated and updated by a tool that scans through the Mentor sites. The user selects one machine, material, and location, then hits the “Send to Printer!” button to send an email message to the selected machine’s operator.Figure 18: AM-Select results screen.Figure 19: AM-Request user interface.4.4IDMI Design-Manufacturing Training MaterialsA wide variety of learning materials has been developed and assembled to support the MENTOR project that spans simple CAD, CATIA V6, systems engineering, materials, manufacturing, and technologies related to the current prize challenge.

The scope of these materials is shown in Figure 20, which shows the main MENTOR learning navigation page. Clicking on an icon launches a web page showing learning material, brings up a set of more specific icons, or launches a software application. For example, the set of icons shown in Figure 14 is displayed to the user when they select the “Design-To-Manufacture Transition” icon shown along the top row.As part of the learning materials, a set of tutorials was developed for manufacturing and 3D printing. Manufacturing processes were classified as subtractive, forming, assembly, and additive and descriptions of each were provided. In additive manufacturing, overall concepts and process descriptions were provided, along with descriptions of specific processes, including fused-deposition modeling, stereolithography, laser sintering, and ink-jet printing.

Links to external resources, including videos, have been provided.A materials tutorial was completed that classifies materials into metal, plastics, ceramics, and composites. Descriptions of each class were included. An introduction to material and mechanical properties was incorporated in the materials so that relationships to 3D printing materials and their properties could be made.Tutorials on the design-to-manufacture software components have been created, specifically for AM-Select and AM-Request. A design for additive manufacturing guidebook will be provided by the end of the project.Figure 20: MENTOR learning materials main web page.5 ClosureThe DARPA MENTOR program is the outreach part of its Adaptive Vehicle Make suite of programs.

Georgia Tech has proposed a plan to involve up to 1000 high schools after 4 years in a series of prize challenges in which students learn 3D CAD, 3D printing, design for manufacturing, some systems engineering, and distance collaboration and teaming methods. A web-based collaborative design-manufacturing infrastructure was presented was developed that integrates CAD, CAE, design-for-manufacturing, and CAM software tools with a network of 3D printers and other manufacturing resources. In distributed teams, students will design, fabricate, and construct electro-mechanical systems (e.g., ground vehicle robots) to perform complex tasks to satisfy the requirements of prize challenges. Many parts they design will be fabricated on 3D printers that are located in high schools or nearby sites.At present, 15 high schools are development partner sites and another 40-50 are engaged to become involved during year 2 of the project.

By partnering with SREB, the AM and systems engineering content is poised to be incorporated into high school curricula, which could have a tremendous impact in deploying content and involving many high schools in Georgia and across the US in the MENTOR program. Educational materials have been developed and assembled to support the MENTOR program and are available on the web, along with software tools for collaboration, design, design for manufacturing, and 3D printing. Through this project, high school students should become excited about pursuing technology-based careers.AcknowledgementThe authors gratefully acknowledge DARPA support through the MENTOR program and the participation of corporate partners Dassault Systemes, 3D Systems, Boeing, and Stratasys. They also thank the many other Georgia Tech project participants for their contributions to the project.References 1.

They also thank the many other Georgia Tech project participants for their contributions to the project.References 1. Thames, J.L., Abler, R., Hyder, A., Wellman, R., Schaefer, D., “Architectures and Design Methodologies for Scalable and Sustainable Remote Laboratory Infrastructures,” Internet Accessible Remote Laboratories, Chapter 13, Azad, AKM, Auer, ME, Harward, VJ (Eds.), Engineering Science Reference, Hershey, PA, 2012.2. Schaefer, D., Thames, J.L., Wellman, R., Wu, D., Yim, S., Rosen, D.W., “Distributed Collaborative Design and Manufacture in the Cloud - Motivation, Infrastructure, and Education,” ASEE Annual Conference, paper AC 2012-3017, San Antonio, 2012.3. Wu, D., Thames, J.L., Rosen, D.W., Schaefer, D., “Towards a Cloud-Based Design and Manufacturing Paradigm: Looking Backward, Looking Forward,” ASME Computers & Information in Engineering Conference, paper #DETC2012-70780, Chicago, August 12-15, 2012.4.

Yim, S., Rosen, D.W., “Build Time and Cost Models for Additive Manufacturing Process in Engineering Conference, paper Information Selection,” ASME Computers & #DETC2012-70940, Chicago, August 12-15, 2012..

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperGUIDED MANUAL DESIGN FOR ADDITIVE MANUFACTURING OF TOPOLOGICALLY OPTIMIZED LEGACY TOOLING PARTSHannah Luben*, Nicholas Meisel† * Department of Mechanical Engineering, The Pennsylvania State University, State College, PA 16801 † School of Engineering Design and Innovation, The Pennsylvania State University, State College, PA 16801AbstractDesign for Additive Manufacturing (DfAM) is a unique conceptual way to adapt a part for Additive Manufacturing (AM). While some of the choices made in DfAM become second nature to seasoned AM designers, inexperienced designers may not know the nuances involved in what is still a developing manufacturing technology. Topology Optimization (TO) in particular tends to create organic shapes that may not be immediately conducive to printing through AM.

Topology Optimization (TO) in particular tends to create organic shapes that may not be immediately conducive to printing through AM. This paper proposes a comprehensive workflow tool to guide a designer, regardless of their experience, through the decision-making process inherent to DfAM. The guide helps the designer manually edit a legacy tooling design into a topologically optimized part that is readily manufacturable through AM. Discussion of a relevant case study follows the outline of the design tool to exemplify its use.1. IntroductionAM is becoming more commonplace as an end-use manufacturing method and is often used in conjunction with topology optimization (TO) [1]. TO is a method of removing nonessential material from a physical part. This weight reduction adds value to the part by more efficiently leveraging material when compared with traditionally designed counterparts.

This weight reduction adds value to the part by more efficiently leveraging material when compared with traditionally designed counterparts. The TO simulation’s constraints are generally the loading conditions the part will be put under and any areas within the part that are manually excluded from the design space of the simulation [2]. The resultant geometry produced by the TO simulation typically is organic and smooth, which is often different from the blocky, simple geometries common in many traditionally manufactured parts. This is part of the reason why it is a popular pairing to design a TO part and then physically create it with AM. Using TO to reduce the weight of a part adds value to that part as per the E3 product value ideology. The E3 concept categorizes a product’s value into its economical, ecological, and experience impacts [3]. Weight reduction from TO adds ecological value, while the small batch sizes and minimal waste material intrinsic to AM add economic value.

Weight reduction from TO adds ecological value, while the small batch sizes and minimal waste material intrinsic to AM add economic value. These added values are especially important in the context of legacy tooling parts; a lighter tool often means less fatigue for the end-user who holds it and other parts it may interface with [4].Despite the material reduction benefits inherent to the use of TO, conventional TO simulations typically create parts “blindly,” without consideration for additive manufacturability. Therefore, some modifications may have to be made to a TO part for it to print as efficiently as possible. Laser-based powder bed fusion (PBF), in particular, has specialized manufacturing concerns, such as consideration of downskin angles, inclusion of drainage holes, and recoater friction [5].

Laser-based powder bed fusion (PBF), in particular, has specialized manufacturing concerns, such as consideration of downskin angles, inclusion of drainage holes, and recoater friction [5]. Though complex, AM-specialized TO software does exist, entry-level designers may not have access to it.1923As such, the goal of this research is to aid a designer, no matter their level of experience, in making manual modifications to geometry that has been generated with a simple, conventional TO simulation. For this paper, the focus is specifically on tooling parts. They are typically small-batch applications since they will generally be replacing lost or broken legacy parts from an established inventory of tools. Small batch applications are typically far more favorable for AM than for traditional manufacturing methods since the latter often has a high up-front cost [6].1.1.

Small batch applications are typically far more favorable for AM than for traditional manufacturing methods since the latter often has a high up-front cost [6].1.1. Establishing Guidance for Design Through AM There are many different DfAM tools that have been created to guide a designer through creating or modifying a part to be more readily manufacturable through AM [7]. Rosen et al. created a DfAM system that provides an overview of the process from start to finish. It comprises the design, CAD, and manufacturing stages, which are then further broken up into subprocesses [8]. In the context of this research, the TO simulation is interchangeable with the design stage since it generates the overall structure of the part, and the manual DfAM operations take place in adjusting the CAD model for the purpose of manufacturing. Ponche et al.

Ponche et al. further broke down the DfAM procedure into optimizing the part orientation, functionality, and toolpaths, which is the basis for the procedure laid out in Section 2 [9]. The focus of this research is the detailed stages of altering the orientation of the part and finetuning its macrostructure.Extensive guides exist for designers who begin the part creation process from scratch with AM in mind [10], but similar principles can be applied to the act of editing a part whose general structure has already been determined. Designing or redesigning a part can be either process-driven or designer-driven, where the former is dictated by a program and the latter is manually done by a person [11]. Specific design considerations must be made in both cases: either indirectly by adjusting settings within the design program or directly through the designer’s active decision- making.

Specific design considerations must be made in both cases: either indirectly by adjusting settings within the design program or directly through the designer’s active decision- making. Selecting the optimal orientation of the part on the build plate is consistently called out as being a vital part of the pre-printing process [6, 9, 11, 12]. The part orientation affects support material use, feature resolution, surface finish, and the anisotropy of the part due to layering [6]. Some of these attributes that cannot be adjusted further through reorientation may be achievable through slight redesigns of the part [11].1.2. DfAM Principles Considering Topology Optimization Topology optimization algorithms can be modified to accommodate certain features of AM, but that is not the focus of this research. To do so is mathematically intense [13, 14, 15], and strays too far from the aim of assisting a designer in modifying a TO geometry created by as simple a simulation as possible.

To do so is mathematically intense [13, 14, 15], and strays too far from the aim of assisting a designer in modifying a TO geometry created by as simple a simulation as possible. For example, Mirzendehdel and Suresh presented a TO algorithm that attempts to minimize support material through various means as a secondary objective, but it required extensive manipulation of the programming behind the simulation solution method (Pareto level-set, in this case) [16]. Even when TO algorithms are modified to work in conjunction with AM, they still require design adjustments before they are ready/optimal to print [17, 18]. Lindemann et al. suggested several ways to go about doing this, one of which was to take the TO simulation output, apply DfAM principles to it, and then model those alterations with standard CAD features [19]. However, the details of applying DfAM principles were not laid out explicitly or robustly.

However, the details of applying DfAM principles were not laid out explicitly or robustly. Though some commercial software can merge TO and DfAM, these licenses can be very costly [20], especially in addition to CAD or slicing software, which are typically the respective preceding and follow-up steps to a TO simulation. An expensive license may not be1924worthwhile for a designer to purchase if these tooling redesigns are for small-batch, customized parts. AM-specialized programs can also be limited in terms of which printers they support, which is why this research focuses on manual editing of a part that can be customized for any printer. Leary et al. were able to do just that and leverage DfAM principles to create an entirely support- free design from a standard density-based TO simulation [21].

Leary et al. were able to do just that and leverage DfAM principles to create an entirely support- free design from a standard density-based TO simulation [21]. The main design steps chosen for the L-PBF-based DfAM procedure are optimizing the support use, internal voids, surface finish, and fine feature resolution, based on ASTM 52911-1, the standard for designing parts for L-PBF of metals [12]. Similar DfAM guides have been created that list out these design considerations but do not have the detail to navigate and assist an inexperienced designer through making these decisions for a specific part [22, 23].2. Proposed Guide for Legacy Tooling RedesignConsidering these previous methods for integrating TO and DfAM frameworks, a novel design guide is proposed specifically for entry-level AM designers to modify a TO version of a legacy tooling part.

The guide shows the logical progression of major phases (TO simulation and DfAM modifications) between three parts—the legacy design, raw TO part, and DfAM-modified part. The steps described in each phase are intended to navigate a designer through making informed decisions in each major step of the process. The resulting flowchart, Figure 1, is broken up into two general phases: the topology optimization simulation and the DfAM decision-making guide. After the designer accepts the results of the TO simulation, it must be re-imported into a manual CAD environment so the designer can edit the raw TO part to make it more readily printable.In order of precedence, the main four design considerations are (1) supports, (2) closed internal voids, (3) surface finish, and (4) fine feature resolution. After each guided decision has been made, the chronological flow loops back to start at the beginning.

After each guided decision has been made, the chronological flow loops back to start at the beginning. This ensures the process iteratively analyzes every design consideration and doesn’t contradict early decisions later downstream. After any changes have been made, as suggested by the yellow ovals in the figure, the likely effects of these changes are listed. For example, reorienting a part will change the support volume, where this change is symbolized by a Greek delta Δ. This change may be positive or negative, so it is shown in yellow. There are certain part redesign actions which may result in less mass or more stiffness, both of which are desirable and colored green. Conversely, negative effects, such as an increase in mass or decreased stiffness, are colored red. These effects are based on whether the designer is adding or removing material to the part. Removal of material may make a part slightly weaker, but also decrease the overall part weight.

These effects are based on whether the designer is adding or removing material to the part. Removal of material may make a part slightly weaker, but also decrease the overall part weight. One such example is widening an internal channel’s cross-section to account for post-processing. This will decrease the stiffness of the part but also its weight. The flowchart references essential surfaces (ES), which are features whose performance will be negatively affected by having a final geometry differing from the original designed CAD part. Examples of this would be screw threads, mating surfaces, or handles gripped by a user. These ES may be part of the TO simulation constraints and they will drive much of the support and surface finish-related considerations in the DfAM portion of the flowchart.2.1.

These ES may be part of the TO simulation constraints and they will drive much of the support and surface finish-related considerations in the DfAM portion of the flowchart.2.1. Topology Optimization Phase Because of the cost and machine limitations of AM-specific TO software, this design tool aims to guide a user through modifying the output of a commercial, non-specialized program. This will reduce up-front costs and solidify a fundamental understanding of DfAM for companies and1925designers who are new to AM. Ideally the chosen simulation software should include CAD modeling capabilities to further reduce costs. The inclusion of TO capabilities is crucial, since one of the desired goals for the redesign of the proposed legacy tooling parts is to make them as light as possible. The TO simulation should include relevant objectives, which are what the simulation will try to achieve, and constraints, which will limit the simulation from achieving the objective [6].

The TO simulation should include relevant objectives, which are what the simulation will try to achieve, and constraints, which will limit the simulation from achieving the objective [6]. Both are necessary components, and therefore require the designer to balance the tradeoffs of prioritizing one over the other [2]. A typical goal is to minimize weight, while constraints may be to retain strength and avoid removing material around, say, a bolt hole. The simulation will take away as much material as it can without interfering with the constraints. A more aggressive mass reduction objective is ideal, when possible, to counteract any material that may be added in the DfAM phase of the process, as will be shown later in Section 3.Figure 1. A simplified version of the entire design guide for legacy tooling redesign. Each step, noted in orange boxes, has its own subtree(s) shown in the corresponding Figures 2 through 5.19262.2.

A simplified version of the entire design guide for legacy tooling redesign. Each step, noted in orange boxes, has its own subtree(s) shown in the corresponding Figures 2 through 5.19262.2. Checking Supports Supports are a necessary component of AM, but inevitably waste material, degrade surface finish, and increase post-processing time after the printing process is complete. Therefore, reducing support material use will typically make the part more easily manufacturable. Some TO programs will have the option to minimize support material prior to generating a design, but this decision framework aims to assist a designer who needs to alter support material manually. As a general rule, surfaces that have an overhang less than 30-45° in the ZX or ZY plane will require support material beneath it to print [12, 24]. TO simulation results often include such overhangs, unless AM rules are incorporated into the algorithm.

TO simulation results often include such overhangs, unless AM rules are incorporated into the algorithm. The three subcomponents in this design step are eliminating support interfaces with ES, reducing the amount of support needed, and ensuring the supports can easily be removed after printing.Eliminate Support-ES Interfaces. The first stage in the DfAM process is to check if support is interfacing with ES. It will be difficult to completely remove these supports after printing, which will disrupt the functionality of the ES. Since these surfaces cannot be redesigned, the only option is to reorient the entire part [12]. The designer should check to ensure this reorientation does not place the part out of bounds of the print volume [19]. If it does, a different reorientation should be attempted.

The designer should check to ensure this reorientation does not place the part out of bounds of the print volume [19]. If it does, a different reorientation should be attempted. In cases where support cannot be eliminated from the ES, a designer should aim to remove as much support material as possible via reorientation in this iterative manner.Reducing General Support Use. Now that supports on the ES have been taken care of, the designer can begin reducing the overall volume of support material across the entire part. Reorientation is typically the quickest, most intuitive route to save material, since it does not involve an extensive redesign of the part [12]. Reorientations should aim to maximize the number of surfaces that are at or above a 45° angle with the build plate, since these will not require support material.

Reorientations should aim to maximize the number of surfaces that are at or above a 45° angle with the build plate, since these will not require support material. After reorienting, it is important to check that the part remains inside the build volume and that the reorientation did not cause supports to interfere with any ES. If the newly oriented part cannot fit in the build volume or it now has support-ES interfaces, a new orientation should be attempted. The designer should attempt another orientation to further reduce the support material until no more improvement in support volume reduction is seen. A new reorientation would entail checking again if the part is outside the build volume and if supports interfere with an ES.The designer then reaches the decision of whether a part redesign would reduce support use. If the part cannot be redesigned, the designer should follow the arrow that exits Step 2 and returns to the main four steps in the DfAM portion of the flowchart.

If the part cannot be redesigned, the designer should follow the arrow that exits Step 2 and returns to the main four steps in the DfAM portion of the flowchart. If these overhangs can be redesigned, the designer should decide between either adding or removing material in the following ways: the angle of the surface in the ZX or ZY plane can either be adjusted so it is greater than the minimum self-supporting angle or the material beneath it can be filled in. Increasing the angle of the overhang will generally reduce both the mass of the final part and its stiffness, while filling in the area beneath the overhang will increase both. While reducing the part mass would better fit with the overall goal of the TO of the part, it may not always be possible, which is why the designer is presented with this choice.1927Figure 2. Fully detailed sub-trees for Step 2: Support Removal.Facilitating Support Removal.

Fully detailed sub-trees for Step 2: Support Removal.Facilitating Support Removal. Being able to remove supports from a part is crucial as they are intended to be temporary and may interfere with part performance if left in place. The general concepts in this section are to avoid support use internally, where it would be difficult or impossible to access them, and to ensure external support is accessible to support removal tools.1928Hollow geometries in the ZX or ZY plane, such as internal channels or holes, can be difficult to access after the part has been built. However, they can be redesigned to have diamond or teardrop shaped cross-sections, which don’t require support material [12, 24]. If the hollow geometry cannot be redesigned to have an alternative shape for the cross-section, the entire part should be rotated so the cross-section of the hollow geometry is parallel to the build plate.

Once again, any reorienting should include a check that the part does not exceed the build volume and does not add support material to ES. If it does either of these, a different orientation should be attempted. The hollow geometry should also be wide enough to post-process for an improved surface finish, i.e., with bead blasting or a small file, if that is necessary for the part to function (i.e., fluid flow). If the hollow geometry is an ES, allowing post-process access is vital. If the channel is too narrow to accommodate post-process tools, it should be widened. External features should also be far enough apart that support material between them can be removed; this will depend on how the supports are intended to be removed after printing. For example, needle-nose pliers will not need as much space to remove supports as regular pliers. The external features should first be redesigned, if possible, to accommodate post-processing.

For example, needle-nose pliers will not need as much space to remove supports as regular pliers. The external features should first be redesigned, if possible, to accommodate post-processing. If the part cannot be redesigned, the designer should attempt rotating if it will reduce support in the region in question. However, rotation should be a last resort by this point to avoid adding in support elsewhere, as this stage should already have minimized support use across the entire part and ES.2.3. Checking Internal Voids In L-PBF, it is important to avoid voids that have no outlet to the external environment surrounding a part. Since the powder feedstock is present in every layer and only solidifies where the final part will be, powder will get trapped inside voids that are completely enclosed [12]. The way to avoid this is to put in drainage holes, so the powder can be poured and vacuumed out of the part once the build has finished.

The way to avoid this is to put in drainage holes, so the powder can be poured and vacuumed out of the part once the build has finished. So, if the part has any fully enclosed surfaces, drainage holes should be added to allow the powder to escape. The drainage holes should be designed with a cross-section that does not interfere with prior decisions, such as internal channels in the ZX or ZY plane requiring diamond or teardrop shapes. The size should be several times larger than the largest particle diameter noted on the powder size distribution (PSD) for the feedstock that will be used to create the part. A suggestion would be five times, or greater, than the maximum of the PSD or the focus diameter of the machine in question, whichever is larger. This ensures that the machine can still create the holes and that powder should pass through them without a problem.Figure 3. Fully detailed sub-tree for Step 3: Checking Internal Voids.2.4.

This ensures that the machine can still create the holes and that powder should pass through them without a problem.Figure 3. Fully detailed sub-tree for Step 3: Checking Internal Voids.2.4. Checking Surface Finish Surfaces that are printed on downward facing overhangs are called downskin, and they typically have a poorer surface finish than upskin surfaces since they meld with some of the loose powder beneath [12]. Therefore, if the part has any ES printed as downskin, the designer should evaluate1929whether this is acceptable [24]. Some factors that go into this decision may be whether the surface finish of these downskin sections can be improved by post-processing and whether enough of the ES is printed with downskin that it will present a problem.Figure 4. Fully detailed sub-tree for Step 4: Checking Surface Finish.1930If only a small portion of the ES will be downskin, it may be acceptable to achieve the desired surface finish with special post-processing.

If it is not, the designer must decide if the part can be redesigned to reduce the amount or increase the angle of downskin. If the part can be redesigned, the designer should attempt one of these methods. If the part cannot be redesigned to reduce the downskin of ES, the designer may try reorienting. However, this should typically be a last resort. Generally, it is more time and cost-efficient to select an orientation that best reduces the support volume rather than prioritizing reducing the downskin in ES; the latter may result in a large volume of support waste and may affect more of the surface finish on the rest of the part. If the designer decides that the downskin of the ES is unacceptable, the part may be reoriented so that the ES is printed as upskin.2.5. Checking Fine Feature Resolution At this point, there should be no further large changes to the part design; this stage should be minimal adjusting of fine features.

Checking Fine Feature Resolution At this point, there should be no further large changes to the part design; this stage should be minimal adjusting of fine features. One of the limitations of any AM process is its feature resolution. For L-PBF, this is reliant upon the laser spot size, which is the diameter of the laser where it contacts the substrate. Commercial printers typically have the minimum feature size, which is the solid feature measurement that can be printed with as few consecutive laser passes as possible, listed as part of its specifications. Negative spaces between adjacent features and the size of positive features need to be considered for the laser resolution to achieve final part accuracy.Figure 5. Fully detailed sub-trees for Step 5: Checking Fine Feature Resolution.1931Adjusting Positive Features. If there are small positive features, there is a chance they may not print [12, 24].

Fully detailed sub-trees for Step 5: Checking Fine Feature Resolution.1931Adjusting Positive Features. If there are small positive features, there is a chance they may not print [12, 24]. If the features are smaller than the prescribed minimum feature size, they must either be eliminated (if deemed unnecessary) or thickened to match the minimum positive feature size. TO simulations often result in symmetric parts, either due to symmetric loading conditions or an intentional design constraint, so features that are asymmetric could potentially be considered for elimination. When stiffness is used as the TO objective, the designer can also observe if these features only connect to the part on one end. In such cases, features that protrude from the part and truncate before connecting somewhere with the part again do not add strength to the part and can be eliminated.Adjusting Negative Features.

In such cases, features that protrude from the part and truncate before connecting somewhere with the part again do not add strength to the part and can be eliminated.Adjusting Negative Features. If the negative spacing between adjacent features is too close, the printed material will sinter together [12, 24]. The features will not be distinct as they were designed to be. This can be avoided by increasing the spacing between adjacent features to be at least the minimum negative feature size. Once again, for symmetric design spaces undergoing symmetric loading, the symmetry of the part can be used as a guide to determine the necessity of these negative spaces. These negative spaces can also be considered unnecessary if the volume they take up is reasonably small. A sample metric for this could be if the negative space is less than or equal to five times the minimum positive feature size.

A sample metric for this could be if the negative space is less than or equal to five times the minimum positive feature size. If the designer decides that the spaces are unnecessary based on such a metric, the negative space can be intentionally filled in with material.2.6. Rotating Part to Minimize Recoater Interaction In L-PBF, the powder feedstock is spread over the build plate by a long, thin recoating mechanism at the start of each layer. Since the layers are on the micron scale, the recoater comes very close to the part on the prior layer and can cause friction if there are any discrepancies in the height of the solid part. Therefore, a common practice to reduce friction with the recoater is to rotate parts around the z-axis, typically so they are approximately at a 45° angle with the recoater blade [12].3. Wrench Case StudyIn this section, the workflow from Section 2 is demonstrated as applied to a simple wrench design.

Wrench Case StudyIn this section, the workflow from Section 2 is demonstrated as applied to a simple wrench design. This legacy part was chosen because of the large amount of solid material that can be used as a design space. It has historically been created with traditional manufacturing methods, but a fast turnaround rate to replace a broken or lost part by quickly printing a new one would be valuable. A weight reduction will improve the ease of use for the end user and ensure a more efficient build. The legacy part has two holes that interface with other parts, which must be excluded from the design space, and since the part is not flat, it will require adjustments to reduce support material that will showcase the decisions made through following the design guide. The original part is shown in Figure 6. The part is fixed at the square end with an assumed torque of 34 Nm applied on the interior of the circular end. A plain alloy steel was used for the sample material.3.1.

The part is fixed at the square end with an assumed torque of 34 Nm applied on the interior of the circular end. A plain alloy steel was used for the sample material.3.1. Topology Optimization of Legacy Design Using SolidWorks 2022, the TO simulation was set up with the goal to reduce the part weight by 85%. The raw output of this simulation is shown on the right side of Figure 6. This aggressive goal was chosen to account for some material that may have to be added back in during the DfAM adjustment phase. The two ES of this part are in the interior of the circular and square cut-outs, which is where the part will interface with other tools. Since these surfaces are assembly interfaces,1932their geometry must be as close to the CAD model design as possible. Note that the middle of the part is not intended to be a handle, and therefore is not considered an ES.Figure 6. Legacy (left) and raw TO (right) wrench designs.3.2.

Note that the middle of the part is not intended to be a handle, and therefore is not considered an ES.Figure 6. Legacy (left) and raw TO (right) wrench designs.3.2. Checking Supports Before beginning the DfAM portion of the flowchart, the raw TO part must be imported into slicing software that has support-generating capabilities. Materialise Magics 26.0 was the program used for this case study, and an EOS M280 profile (including build volume size) was chosen. Note that each orientation/edited part has been placed on the build plate with the software’s “default position” translation operation. This places the part as close as possible to the origin on the build plate and raises it half an inch above the build platform. Supports will be generated to fill the void between the part and the build plate to ensure the part is not damaged during removal post-print. Figure 7 shows the automatic part orientation with non-solid supports shown in blue.

Figure 7 shows the automatic part orientation with non-solid supports shown in blue. The support volume use, 0.1683 in3 in this case, is noted for comparison in potential future orientations.Figure 7. Wrench imported to Materialise Magics in automatically generated orientation. Supports are shown in blue.Eliminate Support-ES Interfaces. Following the flowchart, the first support sub-tree is concerned with supports interfacing with ES. The automatic orientation in Figure 7 places supports inside the circle and square cutouts on either end of the part. These are ES and therefore cannot have support interfaces, so another orientation should be tried. The most apparent orientation to reduce support use in ES is to rotate the part 90° around the x-axis so it stands upright, as in Figure 8. The user should check that the part is still within the build volume and is the appropriate distance from the build plate for removal after printing.

The user should check that the part is still within the build volume and is the appropriate distance from the build plate for removal after printing. This satisfies the ES preservation sub-tree, and the user can move onto the next sub-tree, noting that the support volume has now increased to 0.3512 in3.1933Figure 8. Wrench rotated to eliminate support use interfacing with ES.Reducing General Support Use. The next sub-tree is concerned with general support volume. The second orientation, while eliminating the amount of support material that interfaces with ES, has increased the overall support volume. The first question is whether the part can be reoriented without interfering with the goals of prior reorientations. If so, multiple reorientations should be attempted until the support use has been minimized. Figure 9 shows the result of these iterations, where the part is now angled on the build plate.

If so, multiple reorientations should be attempted until the support use has been minimized. Figure 9 shows the result of these iterations, where the part is now angled on the build plate. Each reorientation should be checked to ensure it still fits within the build volume of the machine. The reorientation found to be the most helpful was a rotation of 50° around the x-axis from the original raw import. The support volume is now 0.1032 in3, which is a 70.62% reduction of material from the second reorientation.Figure 9. Wrench rotated to reduce support use across the entire part and on ES.Now that the part has been reoriented as much as possible to reduce support use without interfering with the ES, the designer can attempt to redesign part surfaces to reduce support use. There are small areas where the overhang angles might be altered to reduce the use of support material beneath them. These are numbered in Figure 11A.

There are small areas where the overhang angles might be altered to reduce the use of support material beneath them. These are numbered in Figure 11A. The areas were identified because they use a large, typically tall, amount of support material. These overhang angles could either be increased or decreased, as shown in Figure 10, which displays a closeup of Area 2 in Figure 11A. However, if the angle is decreased, the supporting member that comprises the overhang will be cut off and eliminated, which is undesirable. Major changes to the TO like this could have a negative impact on the strength of the final part. Therefore, the blue arrow in Figure 10 should be chosen to increase the angle of overhang for Area 2, which will add part material to reduce support material. This same scenario presents itself for Areas 1, 3, and 4, and material was added in all four situations to reduce the support use.

This same scenario presents itself for Areas 1, 3, and 4, and material was added in all four situations to reduce the support use. The result of these changes is shown in Figure 11B, which reduced general support use by about 6.5% from the orientation chosen in Figure 9.1934Figure 10. Diagram of potential fixes to improve support use on a member. The angle of the overhang (shown in green) can either be increased (blue) or decreased (red).Figure 11. Wrench geometry after completion of major reorientations from front before (A) and after (B) redesigning to decrease overall support volume. Multiple annotated supports are noted in A that were removed or lessened in B.Facilitating Support Removal. The next sub-tree in the chart is to consider the feasibility of support removal. While there are no hollow geometries in this part, there are many external narrow geometries within the lattice that will need support material to be removed after printing.

While there are no hollow geometries in this part, there are many external narrow geometries within the lattice that will need support material to be removed after printing. Several examples of this are pointed out in Figure 12A. Following the flowchart, the part can be redesigned to improve the use of support material in these locations. The roughly square-shaped cutouts created by the lattice (see Areas 1, 2, and 3 in Figure 12A) can be considered the overhangs, which means that they can either be manipulated so the angle with respect to the build plate is 45° or greater, or the material can be filled in completely. Similar to the changes made in the general support sub-tree, if the angle of the overhang were to increase, it would reduce the member above it to a thickness smaller than the minimum feature size or eliminate it entirely. Since this member was identified as necessary by the TO simulation, removing it is not an option.

Since this member was identified as necessary by the TO simulation, removing it is not an option. Instead, material should be added to fill in the space beneath the overhang, as shown in Figure 12B. Recall that the aggressive weight reduction goal in the original TO simulation allows for adjustments like this to be made. Note that the supports on Area A3 were not eliminated, but the supports that connected that overhang to the rest of the part have been made unnecessary.1935Figure 12. Comparison of support use in internal lattices before (A) and after (B) adding material.3.3. Checking Internal Voids The next step in the flowchart is to ensure internal voids have powder drainage holes. This part has no true internal voids, but the interior of the lattice can be considered such for the purposes of the case study.

This part has no true internal voids, but the interior of the lattice can be considered such for the purposes of the case study. The EOS M280 has a laser focus diameter of 100μm [25], so if a sample PSD ranges from 20-80μm [26], the designer should create drainage holes of diameter 0.5mm, based on the five times rule suggested in Section 2.3. The large cut-outs in the lattice are more than sufficient for powder drainage purposes, so no drainage holes need to be added to the part.3.4. Checking Surface Finish There are portions of the ES that will be printed with downskin in the current orientation, which can be a concern for surface finish. If this is unacceptable, as the flowchart prompts, the designer would need to choose a different orientation. In this scenario where the downskin is unacceptable for the ES, the designer would return to the second orientation (Figure 8) and then proceed back through the support redesigning stage with that as the permanent orientation.

However, the geometry of the ES in the wrench is very simple and could easily be machined in post-production if the tolerances were extensively disrupted in those areas. Therefore, the current orientation will be retained for the duration of the case study, since the designer is aware of the downskin but has decided it is acceptable not to prioritize.3.5. Checking Fine Feature Resolution The final design step is to consider minimum feature size. A sample laser resolution of 500 microns was used to test minimum feature size, so positive features cannot be smaller than this and negative spaces between features cannot be closer than this.Adjusting Positive Features. The two yellow highlighted regions in Figure 13A show positive features that were measured with the Evaluate tool in Solidworks and found to be much too small for the laser to resolve with a resolution of 100μm. Area 1 is an incomplete strut, which is a common occurrence in TO simulations.

Area 1 is an incomplete strut, which is a common occurrence in TO simulations. This appears to be an inconsequential artifact left over from the simulation. It is not load-bearing, as it does not connect anywhere, and it is not mirrored on the other side. Therefore, it appears to be safe to say that this member is not necessary and can be eliminated. Conversely, the member highlighted in Area 2 is connected to multiple other portions of the part and has a strut that mirrors it on the other side of the part. Thus, this member is a necessary feature, and was consequently thickened.1936Figure 13. Wrench with yellow highlighted areas indicating positive features and green indicating negative features that are too small to resolve before editing (A) and after editing (B).Adjusting Negative Features. The green highlighted area in Figure 13A has a lateral width of about 2.4mm, which is approximately 2.5 times the laser resolution.

The green highlighted area in Figure 13A has a lateral width of about 2.4mm, which is approximately 2.5 times the laser resolution. This area is below five times the minimum feature size threshold, and thus should be adjusted. Since it is close to this value and has thick struts around it, the negative feature should be widened. A similar gap is present on the back of the part, but the width of the second gap is only 0.737 mm wide. This gap is much too small to consider keeping, and thus should be filled.Rotating Part to Minimize Recoater Interaction. Finally, with the major XY reorientations and redesigns completed, the final change to the part must be rotating it 45° about the z-axis. This will be especially important for the top and bottom sections of the part, where there would otherwise be long stretches of the recoater dragging along the part.

This will be especially important for the top and bottom sections of the part, where there would otherwise be long stretches of the recoater dragging along the part. This change is shown in Figure 14, where the recoater moves up and down, spanning the width of the build volume from left to right.Figure 6. Two orientations of the final modified part in the build chamber. 1 is the original orientation after DfAM and 2 has been rotated to minimize friction with the recoater.These changes complete the final part build, in Figure 15. Table 1 shows how each major change to the part affected the volume of itself and of its support. Note that only 3% material has been added to the final part, while the support volume has been reduced by more than 50%. The figure numbers in the first column of the table direct the reader back to the relevant image in this paper.1937Figure 75. Supports necessary for final, DfAM-optimized wrench geometry.Table 1.

Supports necessary for final, DfAM-optimized wrench geometry.Table 1. Changes in part volume and support volume throughout case study.4. ConclusionThough TO is often used in conjunction with AM, there is still a knowledge gap regarding the DfAM of the TO part. This work has established an initial guiding flowchart for a designer to utilize when transforming a legacy tool design into a topologically optimized, ready-to-print part. A case study was presented with a wrench that followed the flowchart from start to finish, elaborating on crucial steps throughout the process. The final part design and orientation reduced the support material use by more than 50%, with a final support volume of 0.0819 in3. The final part itself had a little over 3% of the original material volume added to it, for a final weight reduction of 80% from the legacy part.This guide, while detailed in the areas it does present, is not all-inclusive.

The ASTM standard for L-PBF also recommends selecting a build direction with respect to loading conditions, avoidance of large part surfaces to mitigate warping from thermal stresses and minimizing z-height of the part to reduce build time. This design guide is intended to be an educational introduction to DfAM in the context of manually adapting a TO tool. Therefore, its focus is on facilitating a designer’s fundamental understanding of basic DfAM principles, and operates under the following assumptions: first, that the printer will not fail during the build and second, that the priorities included in the design guide are more important than minimizing the build time. Additionally, the main macrostructure is determined by the output of the TO simulation, so only minor edits can be made.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperHALOGEN-FREE FLAME-RETARDANT POWDER MATERIALS FOR LASER SINTERING: EVALUATION AND PROCESS STABILITY ANALYSISF. Neitzel*, I. Kletetzka*, H.-J. Schmid*Direct Manufacturing Research Center (DMRC) and Particle Technology Group, Paderborn University, GermanyAbstractThe high flammability of components manufactured by laser-based powder bed fusion of plastics (PBF-LB/P) using standard polyamide 12 (PA12) powder still severely restricts their use in industries such as electronics, aviation, and transportation. A key factor for the further establishment of PBF-LB/P is the expansion of the material portfolio with, for example, refreshable and halogen-free flame-retardant (FR) powder materials. Accordingly, various halogen-free FRs are investigated in this work and evaluated with respect to their use in PBF-LB/P.

Accordingly, various halogen-free FRs are investigated in this work and evaluated with respect to their use in PBF-LB/P. First, their decomposition behavior and mode of action are examined. Subsequently, the additives are dry blended with PA12 to investigate properties relevant for PBF-LB/P, such as particle morphology, thermal behavior, and melt viscosity. Afterwards, test specimens for UL94 vertical flame-retardancy tests are produced by processing the dry blends on an EOS P3 PBF-LB/P system. Finally, the process stability of the process-aged powder blends is investigated by again examining thermal behavior and melt viscosity.IntroductionIn the context of additive manufacturing (AM), laser-based powder bed fusion of plastics (PBF-LB/P), here referred to as laser sintering (LS), is one of the most important technologies for processing engineering plastics. In this process, a thin layer of thermoplastic powder is coated on a build platform, preheated and locally melted by a laser.

In this process, a thin layer of thermoplastic powder is coated on a build platform, preheated and locally melted by a laser. Afterwards the platform is lowered by the height of a single layer and the next iteration of the process starts. This cycle is repeated until every layer of the component is finished. Since the development of this process more than 30 years ago, LS has developed into a technology that not only enables functional prototyping but has also recently been increasingly used for industrial series production. Nevertheless, the number of suitable materials is still severely limited, which means that not all customer requirements can be met. An example for this often-conflicting requirements is the flame-retardancy of polymer LS materials, which is required for aviation, medical, electronics and many more industries.

An example for this often-conflicting requirements is the flame-retardancy of polymer LS materials, which is required for aviation, medical, electronics and many more industries. Currently, commercially available flame-retardant (FR) LS powders either contain hazardous halogenated additives or the reusage of process-aged powder is not possible. [1]In the past, many general investigations on the flame-retardancy of polyamide in combination with various FRs have been conducted [2–4]. However, studies focusing on the development of FR materials for LS are limited. A review on the use of various FRs showed that currently only halogen- and phosphorous-based additives are used in AM, while other groups of FR additives have mostly been neglected [5]. In most of these investigations the goal was to increase the flame-retardancy to fulfill the requirements of a better FR classification.

In most of these investigations the goal was to increase the flame-retardancy to fulfill the requirements of a better FR classification. For example,112halogen-free phosphinate-based FRs with different filling ratios were examined and showed that high filling ratios of 25 wt% were needed to achieve a UL-94 rating of V-1. However, this high filler content leads to limitations in processing and to a reduction in the mechanical part properties [6]. Besides this difficult compromise of flame-retardancy and good mechanical properties, the high amount of FR PA12 waste powder is especially problematic. Typically, only a small portion of powder is used to fabricate parts. According to the material supplier EOS GmbH - Electro Optical Systems (EOS) and 3D Systems, only the use of virgin powder is recommended, as components made from used powder do not have the same FR properties.

Therefore, all excess FR powder cannot be used for further sintering processes and is usually disposed of after one process iteration [7, 8]. Regarding recyclability it was recently found that aged PA12 powder with 15 wt% of organic phosphinates in combination with 1 wt% of nanoclay achieved a UL94 standard rating of V-0 (3 mm), while enhancing tensile and impact properties [9]. Whereby, the process stability of the FRs for a further reusage of process aged powder was not part of the investigation. To develop a recyclable FR LS material in the future, suitable FR additives need to be determined. The aim of this work is therefore the investigation and evaluation of various halogen-free FRs in combination with a PA12 matrix for the usage in LS, while taking the process stability into account.To assess whether a particular FR additive is suitable for LS, general requirements, or key properties for LS powder materials must first be set.

Based on these key properties, selection criteria for FRs are then defined before a selection of technically suitable agents is made. Basically, the most important properties of LS polymers can be divided into extrinsic and intrinsic properties. The former includes particle size, particle morphology and powder behavior. The latter includes optical, rheological, and thermal properties. Apart from this, the synthesis and basic structure of the polymer is also one of the key properties but is not part of the investigations within this work since the focus is on adding FRs to a commercially available polymer material. The following Table 1 provides an overview of the key properties of polymers for LS.

The following Table 1 provides an overview of the key properties of polymers for LS. [10]Table 1: Overview of the key properties for LS polymers [10]ExtrinsicIntrinsicPropertiesSpecificsPropertiesSpecificsParticle morphologyPowder behaviorParticle shape Surface properties Production method Particle size distribution Powder rheologyRheological properties Melt viscosityThermal propertiesOptical propertiesSurface tension Melting point Crystallization Thermal conductivity Heat capacity Absorption ReflectionSince not all material requirements are equally important in an additive application, this work focuses on the powder behavior, rheological, and thermal properties of different FR powder blends. As there is a wide variety of FR additives available for polymers, it is necessary to establish selection criteria for decision making, based on the key properties for LS.

As there is a wide variety of FR additives available for polymers, it is necessary to establish selection criteria for decision making, based on the key properties for LS. These defined criteria are as follows:113The FR composition must only contain halogen-free additives due to restrictions on the use of halogens in e.g., electronic components due to their persistence, bioaccumulation, and toxicity in the environment. Since the base material used in this work is a PA12, the additive should be as compatible as possible. The selected additives should have synergistic effects with other FRs to increase the effectiveness and reduce the amount of additive required. [11, 12]The decomposition temperature of the FRs must be higher than the temperature of the build chamber (approx. 178 °C) and the peak melt pool (approx. 270 °C [13]) to maintain the FR mechanism. In addition, the FR mechanism should start before the polymer decomposes (approx.

178 °C) and the peak melt pool (approx. 270 °C [13]) to maintain the FR mechanism. In addition, the FR mechanism should start before the polymer decomposes (approx. 360 °C for PA12) and, if possible, extend over the entire decomposition area.The particle size of each FR should be larger than 5 µm to avoid cohesive flow behavior and to ensure good coating results during LS as well as good dry mixability. If the FR additive is incorporated into the polymer particle for example by melt mixing, smaller particles could be used as well. Due to the maximum layer thickness of the EOS P3 machines to be used, the particle size must be less than 180 µm, to avoid scoring during coating. For the most efficient process regarding process duration and part quality a layer height of 120 µm is preferred.MaterialsSince the LS market is still dominated by PA12 materials the intended approach is to produce FR materials by dry blending various combinations of FRs with PA12 powder.

As a matrix material a PA12 powder (INFINAM PA 6002 P by Evonik) with improved reusability was selected. The powder exhibits a distinct processing window between the onset of melting (melting point after initial heating at 186.5 °C) and crystallization (maximum peak at 142.5 °C). The volumetric median particle size of the material is 58 µm. For the conducted investigations 100 % virgin powder was used.Based on the criteria established in the previous section and with the aim of covering each relevant group of halogen-free FRs as much as possible, the following additives were selected within the scope of this work [5, 14]:Aluminum hydroxide (Apyral 40CD – Nabaltec)Apyral 40CD is aluminum hydroxide, also called aluminum trihydrate (ATH), and is manufactured by Nabaltec AG. When the temperature threshold of approx. 240 °C is reached, ATH decomposes into aluminium oxide and water.

When the temperature threshold of approx. 240 °C is reached, ATH decomposes into aluminium oxide and water. During this endothermic reaction, energy is consumed from the ignition source and at the same time the aluminum oxide forms a protective layer. According to the manufacturer, the specific density is 2.4 g/cm³ and 90 % of the particles are smaller than 3.3 µm. To meet the requirements for a UL94-V0 rating, high fill rates of about 50 wt% are usually required. [15, 16]Boehmite (Actilox B30 – Nabaltec)Nabaltec AG is also the manufacturer of Actilox B30, an aluminum oxide hydroxide also known as aluminum monohydrate (AOH) or boehmite. The FR mechanism is comparable to that of ATH, but AOH contains less crystal-bound water, which in turn leads to a significant increase in temperature stability. Thus, the decomposition of AOH starts at 340 °C.

Thus, the decomposition of AOH starts at 340 °C. Due to its lower energy consumption during decomposition, boehmite is often used as a synergist with, for example,114phosphorus or nitrogen systems. Nabaltec states that the specific density is 3 g/cm³ and 50 % of the particles are smaller than 2.3 µm. [15]Magnesium hydroxide (Versamag – Martin Marietta)Under the trade name Versamag, the Martin Marietta company sells magnesium hydroxide, which, like boehmite, also releases water through endothermic decomposition at about 330 °C. This produces a cooling effect, reduces oxygen availability, and suppresses smoke formation. The endothermic reaction also produces magnesium oxide ash, which forms a protective layer. The specific density is approx. 2.36 g/cm³ and the particle size is in a range of 2 µm – 4 µm.

The endothermic reaction also produces magnesium oxide ash, which forms a protective layer. The specific density is approx. 2.36 g/cm³ and the particle size is in a range of 2 µm – 4 µm. [17]Organic phosphinate (Exolit OP1400 – Clariant)Exolit OP1400 is a synergistic FR blend produced by Clariant with aluminum diethyl phosphinate (Depal) as the main component and another unknown phosphorus component. Due to the P-P synergy, a significant charring effect is observed with polyamides in flame tests. However, it is likely that a gas-phase mechanism of action is also present. OP1400 was developed for use in polyamides and is therefore temperature stable up to 340 °C. 90 % of the particles of the additive are smaller than 53 µm and the specific density is approx. 1.45 g/cm³.

90 % of the particles of the additive are smaller than 53 µm and the specific density is approx. 1.45 g/cm³. [18, 19]Melamine cyanurate (Aflammit PMN525 – THOR / Nord-Min MC-25J – Nordmann)Since melamine cyanurate is used as a standard for unfilled polyamide plastics, the two FRs to be investigated in this project are Aflammit PMN525 from THOR and Nord-Min MC-25J from Nordmann. Analogous to the phosphorus-containing systems, these two additives also form a protective layer by charring. In addition, both have extremely low water solubility, are thermostable up to approx. 300 °C and have a specific density of 1.7 g/cm³. The particle size of PMN525 (98 % smaller than 8 µm) is significantly smaller (98 % smaller than 25 µm) compared to MC-25J. [20, 21]Zinc borate (Aflammit PCI511 – THOR)THOR also sells a zinc borate called Aflammit PCI511.

[20, 21]Zinc borate (Aflammit PCI511 – THOR)THOR also sells a zinc borate called Aflammit PCI511. This is mainly used as a synergist for intumescent systems, where PCI511 improves the structure of the protective layer and can act as an anti-drip agent. The additive is temperature stable up to approx. 300 °C and has a specific density of 0.28 g/cm³. The size of 50 % of the particles is below 10 µm. [22]Borosilicate glass foam (TROVOpowder B – Trovotech)Trovotech has developed TROVOpowder B, a material consisting of an amorphous borosilicate glass powder. This is to be used primarily as a synergist in nitrogen as well as phosphorus FRs and, when exposed to flame, to support the formation of the intumescent layer by vitrification of the polymer. Processing is said to be possible up to 280 °C and the size of 90 % of the particles is less than 9 µm. The manufacturer does not provide more detailed information on the specific density.

Processing is said to be possible up to 280 °C and the size of 90 % of the particles is less than 9 µm. The manufacturer does not provide more detailed information on the specific density. [23, 24]Expandable graphite (GHL PX90-1HT – Luh)GHL PX90-2 is an expandable graphite produced by the Luh company. When a temperature threshold is reached, blowing agents cause the individual graphite layers of the FR to expand and form a voluminous and thermally stable char layer. Typically, expandable graphite is used as a synergist to stabilize the substrate and reduce flammability. The fire protection properties of the additive depend to a large extent on the particle size, as the protective layer formed also decreases115with decreasing particle size. In this case, 75 % of the particles are smaller than 150 µm, which is also the finest variant in the product range. The manufacturer states that at approx. 230 °C – 250 °C the expansion process of the graphite begins.

The manufacturer states that at approx. 230 °C – 250 °C the expansion process of the graphite begins. The specific density is not known. [25, 26]Typically, as the particle size decreases, the importance of van der Waals forces increases, leading to increased formation of agglomerates and thus reduced powder flowability. These agglomerates cannot be completely broken up during mixing with a drum ring mixer, which can lead to an inhomogeneous distribution of the additives and therefore reduced flame-retardancy. For this reason, a high intensity laboratory mixer (EL5 by EIRICH) was used. Each dry blend consisting of the PA12 matrix material and one or more of the previously introduced FRs was mixed at 500 rpm for 10 min. [10]Experimental MethodsThe decomposition behavior of the various FRs was investigated by thermogravimetric analysis (TGA). The analyses were carried out with an TGA/DSC 1 (Star System) of the company Mettler-Toledo AG according to the DIN EN ISO 11358-1 standard.

The analyses were carried out with an TGA/DSC 1 (Star System) of the company Mettler-Toledo AG according to the DIN EN ISO 11358-1 standard. The specimens with a sample mass of approx. 30 mg were heated with a set rate of 10 K/min under nitrogen atmosphere up to 600 °C. The specimen chamber is then flooded with oxygen and heated up to 1000 °C to burn out residual materials and impurities. During the analysis the change in mass is measured as a function of temperature. [27]Particle size distribution (PSD) was measured by dynamic image analysis using a QICPIC by Sympatec according to ISO 13322-2. Within the dynamic image analysis, a digital camera captures the particle collective flowing continuously through the measurement volume and can thus record the actual shape and size distribution. The measuring range was set to M5, therefore particle sizes in between 1.8 and 3755 µm were measurable. A dry dispersion unit was used to break up the particle collective.

The measuring range was set to M5, therefore particle sizes in between 1.8 and 3755 µm were measurable. A dry dispersion unit was used to break up the particle collective. [28, 29]To investigate the melt viscosity at a given shear rate, the melt volume flow rate (MVR) was determined using a capillary rheometer from Zwick Roell. The test series was carried out in accordance with DIN EN ISO 1133-1. For this purpose, a sample mass of approx. 4 g of the material to be tested was filled into a heated cylinder. The specimen was melted for 5 min. The melt was then extruded through a fixed diameter die by a piston with a defined test weight. To measure MVR, the length of the distance traveled by the piston in a given time, or the time required for the piston to travel a given distance, was recorded, and used to calculate the extrusion rate in cm³/10 min. The test temperature was set at 220 °C with a test weight of 10 kg for all tests.

The test temperature was set at 220 °C with a test weight of 10 kg for all tests. To avoid falsification of the measurement results due to moisture in the polymer, the samples were first dried in a laboratory oven at 105 °C for 30 min. The MVR was calculated from the average of at least three samples per material tested. [30]The analysis of the thermal behavior especially the processing window of each dry blend was carried out using differential scanning calorimetry (DSC). Within the scope of this work, the measurement according to DIN EN ISO 11357-1 was carried out with the DSC 214 Polyma measuring instrument from Netzsch. During the measurement, the sample and a reference are subjected to the same temperature program. The difference in the heat flows is measured, which provides information about the melting and crystallization behavior of the plastic, for example. In116these investigations, the samples were heated to 230 °C at a heating and cooling rate of 10 K/min.

In116these investigations, the samples were heated to 230 °C at a heating and cooling rate of 10 K/min. The sample mass was 10±0.3 mg. [31]Basically, a fire protection classification of plastics can be carried out according to the UL94 regulations of the American Underwriters Laboratories. The test was carried out on a test bars (80 x 12.5 x 3 mm) with a methane gas flame of 50 W and a flame length of 20 mm. In the vertical fire test, the test specimen was flamed twice for 10 s each. After each flame treatment, the burning duration, and the dripping of burning material were evaluated with the aid of a cotton swab. The specimens were pretreated at 70 °C for 7 days in a hot air oven. Depending on the burning time and dripping of the burning material, the classification was made according to the thickness into flammability classes V-0, V-1, V-2, where V-0 is the best classification. Table 2 shows the criteria and classification of the materials rating in the UL94 V test.

Table 2 shows the criteria and classification of the materials rating in the UL94 V test. For each dry blend at least five specimens were tested. [32]Table 2: Criteria for flame-retardancy classification in UL94 vertical tests [32]Criteria for sample classificationV-0V-1V-2Burning duration for each individual specimen after first or second ignition Total burning duration for any set of five specimens after both ignitions Glow duration after for each individual specimen after second ignition Complete burnup of the specimen until clamping Ignition of cotton wool indicator by burning dripping of the specimen≤ 10 s ≤ 50 s ≤ 30 s No No≤ 30 s ≤ 250 s ≤ 60 s No No≤ 30 s ≤ 250 s ≤ 60 s No YesResultsIn this work the basic procedure for evaluating FRs regarding a suitable LS application as well as process stability is divided into five steps as shown in Figure 1.

After each step the FRs respectively the dry blends are evaluated to determine if the step specific requirements are met. Specimens which don’t fulfill the requirements won’t be included in the following examinations.Figure 1: Methodology for evaluation processFirst the decomposition behavior of the selected FRs was studied to determine if the additive can withstand the laser induced temperatures during the LS process. For test specimen production in the following third step the area energy density was set to 0.0288 J/mm² at which the expected melt pool temperature is approx. 270 °C [13]. Accordingly, the decomposition and therefore mode of action of the FRs to be part of further examinations must start at higher temperatures. For a FR to be processable in LS a weight loss of 1 % was set as the limit in this work which should not be exceeded at a temperature lower than the expected melt pool temperature.

For a FR to be processable in LS a weight loss of 1 % was set as the limit in this work which should not be exceeded at a temperature lower than the expected melt pool temperature. The decomposition behavior was investigated by TGA-analysis and the temperatures of each FR at a weight loss of 1 % and 5 % are given in Table 3.These results clearly show that aluminum hydroxide (Apyral 40CD) reached the weight loss limit of 1 % at approx. 234 °C. At 268 °C at least 5 % of the test specimen weight was lost117thus indicating that this additive will most likely decompose to a certain amount during the LS process at the set energy density and consequently won’t be suitable for a future LS-FR-powder application. The borosilicate glass foam (Trovopowder B-K3) was the first additive to reach a weight loss greater than 1 % at 208 °C.

Here, the percentage weight loss is documented as a function of temperature. Analogous to the previous description the decomposition of aluminum hydroxide has already begun before reaching the assumed melt pool temperature. After losing about 30 % of the sample weight due to the release of water 70 % of the sample weight remain as a residue mostly consisting of aluminum oxide. When looking at the behavior of the borosilicate glass foam it is noticeable that this additive has decomposed to a small extent at 270 °C but most of it remains stable until reaching approx. 280 °C. When the decomposition of this additive is done, about 45 % of the sample weight remain as a glass foam residue. Basically, to increase the flame-retardancy of a polymer by using FRs, the decomposition of these additives should preferably start at the same time or just before the decomposition of the polymer to achieve the highest efficiency [33].

This applies to melamine cyanurate (Aflammit PMN525) and the organic phosphinate (Exolit OP1400) and therefore the assumption is made that these FRs will achieve the best flame-retardancy in a PA12 application. After the decomposition of phosphinate a carbonaceous foam remains as residue. The melamine- based FR as well as the virgin PA12 powder (PA 6002 P VP) decomposed almost residue-free. Compared to aluminum hydroxide boehmite (Actilox B30) contains less releasable crystal-bound water which is why the thermal stability is higher and less weight is lost during decomposition. Also, about 85 % of the sample mass remain as aluminum oxide residue. In addition, the thermally induced water release of boehmite starts at approx.

Also, about 85 % of the sample mass remain as aluminum oxide residue. In addition, the thermally induced water release of boehmite starts at approx. 100 °C higher than the polymer decomposition and therefore besides the naturally worse efficiency the flame-retardancy will be additionally less effective compared to melamine- or phosphorous-based FRs in a PA12 application.118Aflammit PMN525020406080100100200300400500600700800Weight Loss / %Temperature / °CTrovopowder B-K3Actilox B30PA 6002 P VPMelt pool temperatureExolit OP1400Apyral 40CDFigure 2: Decomposition behavior of selected FRs in comparison with unfilled PA12 and the LS melt pool temperature determined by TGA under nitrogen (up to 600°C) and oxygen (from 600 °C to 1200 °C) atmosphereAll due to their decomposition behavior suitable FRs for LS were used to produce dry blends as described in the previous section. The composition of each dry blend is shown in Table 4.

The composition of each dry blend is shown in Table 4. In general, the total filling ratio is based on the manufacturer’s recommendation for injection molding to achieve a UL94-V0 classification. If no recommendation was given a total filling ratio of 15 wt% was chosen. Because of the typical application of zinc borate (Aflammit PCI 511) and boehmite (Actilox B30) both additives serve mainly as synergists in this work.

Because of the typical application of zinc borate (Aflammit PCI 511) and boehmite (Actilox B30) both additives serve mainly as synergists in this work. As discussed before boehmite was not used solely as a FR due to the expected less effectiveness.Table 4: Composition, filling ratio and notation of each dry blendFlame-retardant 1Filling ratio / wt%Flame-retardant 2Filling ratio / wt%NotationAflammit PMN 525 Versamag Versamag Aflammit PCI 511 Nordmin MC-25J Nordmin MC-25J Nordmin MC-25J GHL PX 90/-2 Exolit OP1400 Exolit OP1400 Exolit OP1400 Trovopowder B-K315 25 10 15 14 10 11 25 15 10 10 15Aflammit PCI 511Aflammit PCI 511 Actilox B30Aflammit PCI 511 Actilox B3055 45 5PMN 525 Versa Versa / PCI 511 PCI 511 MC-25J MC-25J / PCI 511 MC-25J / B30 PX 90/-2 OP1400 OP1400 / PCI 511 OP1400 / B30 B-K3In the second step the dry blended FR powders were characterized beginning with the particle size distribution.

Figure 3 shows the d10, d50 and d90 percentiles of the measured particle size distributions. In context of LS the particle size has a big impact on the powder flowability, optical properties, surface quality as well as the detail resolution of the produced parts. A high119percentage of fine powder can reduce the powder flowability due to the tendency to form agglomerates and a high percentage of coarse powder can affect the part quality negatively. [10]It is noticeable that the particle size (d90 = 226 µm) of the dry blend containing expandable graphite (PX 90/-2) exceeds the maximum layer height of the LS machine and is therefore not a suitable powder for LS. Because most of the examined FRs were developed as additives for compounding and not for dry blending particles sizes are smaller than 10 µm.

Because most of the examined FRs were developed as additives for compounding and not for dry blending particles sizes are smaller than 10 µm. This leads in comparison with the unfilled PA12 powder (d10 = 38 µm) to a higher percentage of fine powder indicated by the smaller particle size at d10 percentile (e.g., PMN 525 d10 = 8 µm). Therefore, the powder flowability might be an issue during the processing step.

Therefore, the powder flowability might be an issue during the processing step. The median as well as the particle size for 90 % of the sample volume are approximately the same, as these depend significantly on the polymer particles.80.6981.6578.778.579.179.3379.9780.6878.3178.1378.9678.65225.7955.7955.8855.554.2855.0955.8957.0256.2854.9854.555.6557.9773.758.139.039.9210.0711.2911.7512.8713.3115.5420.0321.3338.1839.97020406080100120Particle size / µmd10d90d50Figure 3: Particle sizes of investigated FR dry blends shown as percentiles (d10, d50, d90)In addition to the particle size distribution, rheological properties of the dry blends were investigated by measuring the melt volume rate (MVR). In general, the coalescence of the polymer particles depends on the viscosity of the melt, which can significantly affect the subsequent properties of the components, such as surface roughness.

In general, the coalescence of the polymer particles depends on the viscosity of the melt, which can significantly affect the subsequent properties of the components, such as surface roughness. Factors affecting the viscosity of the entire particle collective include both the average molecular weight of the polymer and the presence of non-meltable particles in the melt. [10]The results of the measurement are given in Figure 4. In general, the melt volume rate of the dry blends is about 140 cm³/10min, which is lower by 50 cm³/10min compared to the unfilled PA12 powder (191 cm³/10min). Since the FRs do not melt at the measurement temperature of 220 °C, they remain as solid particles during the measurement and thus increase the overall viscosity of the melt due to the increase in polymer shear rate. This is also noticeable for the dry blend with the highest filling ratio containing 25 wt% of magnesium hydroxide (Versa), which results in the lowest melt volume rate of approx.

This is also noticeable for the dry blend with the highest filling ratio containing 25 wt% of magnesium hydroxide (Versa), which results in the lowest melt volume rate of approx. 56 cm³/10min. The powder composition with B-120K3, on the other hand, has a significantly higher melt volume rate of about 300 cm³/10min. In addition, brown dots were visible in the plastic strand extruded during the measurement, indicating thermal decomposition of the additive.

In addition, brown dots were visible in the plastic strand extruded during the measurement, indicating thermal decomposition of the additive. For this reason, the borosilicate glass foam will be declared as not suitable for LS, as thermal decomposition is to be expected during laser exposure, which has already been made visible by the results of the TGA.MVR / cm³/10min050100150200250300350Figure 4: Melt volume rate of investigated FR dry blends of 3 specimens each at 220 °C with 10 kgDifferential scanning calorimetry analyses were performed to study the thermal behavior of each dry blend to determine the melting and crystallization ranges as a function of temperature. For each sample, two heating and cooling cycles were performed in direct succession for an examination independently of the thermal history. In addition, the DSC analysis provides a starting point for adjusting the build temperature in the LS process.

In addition, the DSC analysis provides a starting point for adjusting the build temperature in the LS process. [10]Looking at the results in Figure 5, it is noticeable that the crystallization peak of the FR dry blends is already reached at higher temperatures compared to the unfilled PA 6002 P. The solid particles in the melt serve as nucleating agents and thus cause a shift in the crystallization range [32, 34]. Since only the polymer particles melt and the matrix polymer is the same for all dry blends, the melting point remains unchanged at 187 °C. In addition, the sintering window of the dry blends containing melamine cyanurate is about 6 °C smaller than for OP1400, for example.

In addition, the sintering window of the dry blends containing melamine cyanurate is about 6 °C smaller than for OP1400, for example. Possibly, due to the smaller particle size of MC-25J compared to OP1400, there are significantly more solid particles in the melt, which can act as nucleating agents and initiate the crystallization of the polymer.121TcTmTemperature / °C187.0187.0187.0187.0187.0187.5187.5187.5187.0187.0186.5187.0144.0144.5146.0146.0146.5148.5149.0150.5151.0151.5152.0152.075100125150175200225Figure 5: Melt and crystallization temperature of investigated FR dry blendsThe DSC diagram (cf. Figure 6) of the dry blend containing OP1400 also shows that the melting enthalpy of the dry blend is lower, since the filler does not melt. This applies for all analyzed blends.

Figure 6) of the dry blend containing OP1400 also shows that the melting enthalpy of the dry blend is lower, since the filler does not melt. This applies for all analyzed blends. Between the first and second cycle, a clear shift between the melting ranges can also be seen, since the as-received polymer powder has a higher crystallinity than the crystal structure resulting from the DSC with a cooling rate of 10 K/min. The melting point of the FR dry blend and the PA12 powder remains the same, as expected. While for the dry blend the recrystallization peak remains unchanged between the first and second cycle, a shift to lower temperature for the virgin PA 6002 P powder is noticeable. It is assumed that the increasing entanglement and elongation of the polymer chains due to post-condensation during the measurement makes crystallite formation more difficult, which accordingly leads to a decrease in crystallinity. [34]2. OP14002-101234100120140160180200DSC / mW/mgTemperature / °C2.

[34]2. OP14002-101234100120140160180200DSC / mW/mgTemperature / °C2. PA 6002 P VP1. PA 6002 P VP1. OP1400Figure 6: DSC of OP1400 dry blend and neat PA 6002 P powder after first (solid) and second (dashed) temperature cycle; 10 K/min heating/cooling rate122third step by producing flame-retardancy test specimens on a AMCM customized EOS P395 system. To lower the required quantities of powder for one build process the total build volume of the machine was reduced to 100 x 100 x 100 mm while at the same time build chamber temperatures up to 300 °C can be achieved for processing high-performance polymers. The powder is applied with a roll coater.The processability wastheninvestigated withintheFirst, the powder application of each dry blend was visually inspected and evaluated in terms of application behavior. As expected, the quality of the powder bed decreased with increasing fines content, as this strongly affects the flowability of the powder.

As expected, the quality of the powder bed decreased with increasing fines content, as this strongly affects the flowability of the powder. Especially with the PMN525 dry blend, which has the highest fines content due to the smallest particle size (d98 = 8 µm), the application of a homogeneous powder layer without coating defects was not possible (cf. Figure 7 (left)). Although it was still possible to apply a powder layer of sufficient quality for component production due to the installed roll coater, the use of this powder on conventional EOS P3 systems with blade coaters is questionable. This applied for all dry blends with a d10 < 12 µm.

This applied for all dry blends with a d10 < 12 µm. The best coating results were achieved using the dry blends containing OP1400, as these have a much lower fines content because of the significantly larger particles (d90 = 53 µm), as can be seen in Figure 7 (right).Figure 7: Coated powder of FR dry blends with PMN525 (left) and OP1400 (right)Afterwards, the build chamber temperature was determined individually for each dry blend to prevent curling prior to the production of flame-retardancy test bars according to the UL94 test standard. In dependency of the dry blend chamber temperatures vary between 168 °C (OP1400) and 172 °C (Versa). Due to the smaller build volume, the length of the test specimens was reduced to 80 mm, accordingly the geometrical dimensions are 80 x 12.5 x 3 mm. The laser parameters for part production were set to an area energy density of 0.0288 J/mm².

The laser parameters for part production were set to an area energy density of 0.0288 J/mm². In specimens manufactured from powder with PMN525, significant unevenness can be observed in the edge areas of the individual layers (cf. Figure 8 (left)), which is due to the reduced flowability of the powder. As can be seen in Figure 8 (right), the powder with OP1400, on the other hand, was able to produce test specimens without geometrical deviations and surface defects.Figure 8: UL94 V test specimens of FR dry blends with PMN525 (left) and OP1400 (right)In the fourth step, the flame-retardancy of the individual dry blends was examined. Thegoal is to achieve a UL94 V-0 rating to enable a future production of for example electronic123components by LS.

Thegoal is to achieve a UL94 V-0 rating to enable a future production of for example electronic123components by LS. The following Figure 9 shows the total burning duration of five samples of each dry blend and the corresponding flame-retardancy classification.050100150200250Total burning duration / s V-1Figure 9: Total burning duration and flame-retardancy classification according to UL94-V test standard of 5 specimens produced by FR dry blendsThe results show that at a filler content of maximum 10 wt%, no FR achieved a sufficient protective effect, as the dry blends with OP1400 / PCI 511, OP1400 / B30 and Versa / PCI 511 did not reach any classification due to total burning durations > 250 s. The same applies to the two FRs Versamag (25 wt%) and PCI 511 (15 wt%), which also did not offer sufficient fire protection.

The same applies to the two FRs Versamag (25 wt%) and PCI 511 (15 wt%), which also did not offer sufficient fire protection. MC- 25J / PCI 511 is the first dry blend to achieve a flame-retardancy classification, since here the total burning duration of about 221,5 s was below the threshold of 250 s. Due to the strong dripping behavior of the test specimens, which led to the ignition of the cotton wool indicator, this blend was classified in V-2. This can also be seen clearly in Figure 10 (left) from the strand formation in the lower part of the specimens. The test specimens from the melamine cyanurate-containing dry blends MC-25J, PMN 525 and MC-25J / B30 achieved the lowest total burning durations, averaging about 28 s, with the blend containing boehmite as synergist performing best at about 24 s.

Although these are well below the 50 s threshold that would be sufficient for a V-0 rating, burning droplets ignited the cotton wool indicator in all tests (see Figure 10 (middle)), resulting in a V-2 rating. The best rating of V-1 was achieved by the test specimens made from powder with OP1400 (15 wt%). The total burning duration of about 53 s was only 3 s above the limit for a V-0 rating. In addition, the formation of a char layer during the tests, which is typical for phosphorus- containing FRs (see Figure 10 (right)), completely prevented the dripping of material from the heat-affected zone and thus the ignition of the indicator.124Figure 10: Laser sintered UL94 vertical test specimens with MC-25J / PCI 511 (left), MC-25J / B30 (middle) and OP1400 (right)In the last and fifth step, aged powder was collected and re-characterized after the respective LS process to evaluate the process stability of the four dry blends that performed best in the flame-retardancy tests.

For this purpose, first the melt viscosity was investigated by determining the melt volume rate and then the thermal behavior by means of DSC analysis.Typically, the reactive chain ends in PA12 powders tend to post-condense during LS due to the long storage at elevated temperatures near the melting point. In this work, the build process took about 4 h due to the reduced build volume. In industrial use, a process usually takes significantly longer. This long process duration leads to longer polymer chains with a correspondingly higher average molar mass. This leads to an increase in melt viscosity, which is reflected in a significant decrease in the melt volume rate. When processing this post-condensed material, there may be an increased occurrence of surface defects, also known as the orange peel effect, which should be avoided.

When processing this post-condensed material, there may be an increased occurrence of surface defects, also known as the orange peel effect, which should be avoided. [10, 35]PA2210 FR VP020406080100120140160180OP1400PMN 525MC-25J / B30MC-25JMVR / cm³/10minVirgin Powder020406080100120140160180OP1400PMN 525MC-25J / B30MC-25JMVR / cm³/10minAged PowderPA2200 VPFigure 11: Melt volume rate of virgin and aged FR dry blends with best flame-retardancy compared to unfilled and FR PA12 virgin powder of 3 specimens each at 220 °C with 10 kg125The measurement results of the dry blends shown in Figure 11 are compared with unfilled standard PA12 powder (PA 2200 VP) and the commercially available PA12 FR virgin powder (PA 2210 FR VP). In virgin state the dry blends have a higher melt volume rate than PA 2200 VP due to the use of the recycling optimized PA 6002 P as polymer matrix.

In virgin state the dry blends have a higher melt volume rate than PA 2200 VP due to the use of the recycling optimized PA 6002 P as polymer matrix. The effect of polycondensation and the resulting increase in viscosity is visible for all dry blends, as the melt volume rate of the aged powder is reduced by about 35 % on average and therefore lower than the rate of PA 2200 VP. However, the rates of the aged powder are still significantly higher than those of the PA 2210 FR VP. Since no surface defects occurred during the processing of the PA 2210 FR VP in previous investigations, it can be assumed that these will also not occur when using the aged powder of the dry blends.In Figure 12 and Figure 13 the DSC diagrams of both dry blends with the best results in the flame-retardancy tests (OP1400, MC-25J / B30) in virgin and process-aged state are shown. Two measurement runs were performed for each sample.

Two measurement runs were performed for each sample. As expected, the melting point of both materials shifted to lower temperatures (approx. 177 °C) due to the change in crystallinity of the polymer caused by the deviating cooling rate after the first cycle. In comparison between virgin and aged powder the crystallization did not change significantly, so that it can be assumed that the solid particles of the FR were not affected by the LS process. In future investigations with aged material, this assumption must be confirmed by further flame-retardancy tests.2. Virgin Powder1. Virgin Powder1. Aged Powder2-10123100120140160180200DSC / mW/mgTemperature / °C2. Aged PowderFigure 12: DSC-Analysis of PA12 powder with OP1400 in virgin as well as process aged state after first (solid) and second (dashed) temperature cycle; 10 K/min heating/cooling rate1262-101234100120140160180200DSC / mW/mgTemperature / °C2. Virgin Powder1. Aged Powder2. Aged Powder1.

Virgin Powder1. Aged Powder2. Aged Powder1. Virgin PowderFigure 13: DSC-Analysis of PA12 powder with MC-25J / B30 in virgin as well as process aged state after first (solid) and second (dashed) temperature cycle; 10 K/min heating/cooling rateSummaryIn this work, various FRs, e.g., aluminum hydroxide, boehmite, phosphorous and melamine cyanurate, were investigated for their suitability in LS. These additives were dry blended with the commercially available INFINAM PA 6002 P PA12 (optimized for reduced aging effects) powder with total filling ratios ranging from 11 wt% up to 25 wt%. The dry blends were then subjected to a five-step evaluation process, ranging from decomposition behavior and powder characterization to processability during LS, achieved fire protection and process stability.

After each step, the FRs and later the dry blends were evaluated to determine if the step specific requirements were met.As the mode of action of all FRs is initiated by thermal decomposition the decomposition behavior was analyzed first by thermogravimetric analysis. It was shown that aluminum hydroxide is not suitable for LS since a mass loss of > 5 % was measured at the assumed melt pool temperature of 270 °C, therefore the additive would already decompose to a certain amount in the process due to the energy impact of the laser. Measuring the particle size distribution of the dry blends further revealed, that expandable graphite is also not suitable for LS, due to particle sizes larger than the maximum layer height of 180 µm of the EOS P3 system. During the powder characterization, including the measurement of the melt volume rate and a DSC analysis, it was found that the thermal stability of borosilicate glass foam is also not sufficient to be used in LS.

When processing the remaining dry blends, it became clear that especially the powder flowability of zinc borate, melamine cyanurate and magnesium hydroxide due to high fines content caused by the small particle size of the FRs can limit the application behavior in systems with blade coaters. Since the AMCM customized EOS P3 system used in this work has a roll coater, it was still possible to produce flame-retardancy test specimens. However, the reduced powder flowability has in some cases led to significant geometrical deviations of the test specimens. The subsequent fire protection tests carried out in accordance with UL94 V test standards showed that dry blends with a total filling ratio of 15 wt%, with the FR additive with the highest filler content being only 10 wt%, were not sufficient for any classification. Dry blends with melamine cyanurate showed excellent fire127protection properties, but the occurrence of burning droplets resulted only in a V-2 classification.

Dry blends with melamine cyanurate showed excellent fire127protection properties, but the occurrence of burning droplets resulted only in a V-2 classification. The best properties were obtained with the phosphorus-containing FR Exolit OP1400, which achieved a V-1 classification at a wall thickness of 3 mm and a filler content of 15 wt%. In addition, subsequent measurements of melt viscosity and thermal behavior suggest that both the aged powder of the dry blends with melamine cyanurate and with OP1400 may be suitable for further process iterations.In future studies, the process stability of the two most promising FRs, melamine cyanurate and OP1400, must be confirmed by further fire protection tests with a mixture of process-aged and virgin powders. Furthermore, the aim is to achieve a V-0 rating by improving the dripping behavior when using melamine cyanurate and by increasing the filler content of OP1400.

F5 Hierarchies for Additive Manufacturing digital representations and Enhanced AnalyticsL. V. Monnier1, H. Ko3.1University of Burgundy, Dijon, France. 3Arizona State University, Tempe AZ 85281.AbstractAdvancement in Additive Manufacturing (AM) technologies and data acquisition techniques have led to an increase in AM data generated. However, due to the large volume and the diversity of AM data available it is becoming challenging to eﬃciently store, analyze, and represent AM processes. HDF5 has the potential to allow an easy access to big data by oﬀering a hierarchical data catalog. Thus, AM processes could be represented through a hierarchy based on the data analytic needs and directly link the corresponding AM data. This paper investigates the use of data formats to represent big data and AM dataset. Existing AM ontologies and models are reviewed in order to eﬀectively encapsulate AM information and incorporate the hierarchy into an HDF5 AM wrapper.

Existing AM ontologies and models are reviewed in order to eﬀectively encapsulate AM information and incorporate the hierarchy into an HDF5 AM wrapper. Three hierarchies are proposed to represent speciﬁc perspectives of AM processes: the digital twin of AM Product Lifecycle, the AM V model representation, and the material centric characteristics.Introduction Additive Manufacturing (AM), also referred to as 3D Printing, is a manufacturing process that allows the construction of parts layer-by-layer from a 3D model using speciﬁc techniques [1]. Multiple stages have been identiﬁed to represent the AM product lifecycle, the design, process plan, build, post process, test and validation [2]. Each stage generates large amounts of data with various types based on the corresponding stage’s needs. Intensive experiment-based studies can be conducted in order to insure part quality and utilize the full potential of AM processes. However, it can be time consuming and costly [3].

Intensive experiment-based studies can be conducted in order to insure part quality and utilize the full potential of AM processes. However, it can be time consuming and costly [3]. Therefore, computational models and simulation have been developed to help part qualiﬁcation by realizing process digitalization.I.Hierarchical Data Format (HDF5) has been recently applied in diﬀerent domains in order to help manage big datasets. In astronomy, Pourmal et al. uses HDF5 in order to store and quickly access huge astronomical dataset with earth science [4]. NASA has been using HDF5 in order to store images and radio frequency data from their Earth Observing System (EOS) [5]. In Biology, Dougherty et al. uses HDF5 in order to store huge dataset of biological to support multi-dimensional datasets [7]. BioHDF is a HDF5 hierarchy that has been developed in order to represent and store experimental data [8]. In construction, Krijnen et al.

BioHDF is a HDF5 hierarchy that has been developed in order to represent and store experimental data [8]. In construction, Krijnen et al. have worked on translating the IFC standards for buildings in HDF5 [9]. A parser has been developed using the EXPRESS [10] syntax of Industry Foundation Classes (IFC) ﬁles [11] to automatically translate an IFC ﬁle and store it in a HDF5 ﬁle. In Manufacturing, Khan et al. uses HDF5 as a neutral format to allow the transfer of information between Computer Aided Design (CAD) et Computer Aided Engineering (CAE) [12]. Schmitz et al. have introduced Dream3D in order to model and represent in standardized and unique way microstructure [13]. All these diﬀerent domains have proven the beneﬁt and eﬃciency of using HDF5 format in order to easily store and manage big datasets. AM generates big data sets during the diﬀerent processes and such format could help solve AM data challenges.images [6]. Jackson et al.

AM generates big data sets during the diﬀerent processes and such format could help solve AM data challenges.images [6]. Jackson et al. developed an HDF5 modelIn this paper, the goal will be to improve AM process digitalization and data analysis by proposing an approach to help manage the large amounts of data generated from the diﬀerent stages of the AM processes. The paper starts by introducing background work on data management to help organize and eﬀectively improve AM computational eﬀorts. We highlight the AM data challenges and describe the drive to develop better data organization. Then, we propose an approach that uses hierarchical data format in order to optimize data curation1551and enhance data analytics. Finally, we showcase our approach by demonstrating the use of HDF5 as a base for digital twin implementation.II. Hierarchical Organization of AM Data1.

Finally, we showcase our approach by demonstrating the use of HDF5 as a base for digital twin implementation.II. Hierarchical Organization of AM Data1. Data Management to Drive AM Data AnalysisIn order to eﬃciently use and store big datasets, three diﬀerent types of data management techniqueshave been identiﬁed : based on model, based on ontology, and based on hierarchy.Traditional databases are majorly used to maintain and store data in a speciﬁc format using a centralized architecture. Each database is described by a schema based on a model representing the system and the dataset. Lu et al. developed the Additive Manufacturing Integrated Data Model (AMID) in order to capture, store, and properly manage AM data for easy query and analysis [14]. This model has been created using the Product, Process, Resource (PPR) method and is used to represent the data in the Additive Manufacturing Material Database (AMMD).

This model has been created using the Product, Process, Resource (PPR) method and is used to represent the data in the Additive Manufacturing Material Database (AMMD). This traditional method of storing data is a good and proven way to represent datasets but it hardly supports big data, and doesn’t allow for easy transfer of databases. A ﬂat structure is used to describe heterogeneous datasets and doesn’t allow for easy linkage of the data. Thus, the handling of big and complex datasets containing product lifecycle data is challenging and time-consuming. New concepts and data management methods have been investigated in order to bridge these challenges.An ontology, in computer science, is a data model that represents a set of concepts within a domain and the relationships between those concepts. Antoniou et al. deﬁne ontology as “languages that allow users to write explicit, format conceptualizations of domain models” [15].

Antoniou et al. deﬁne ontology as “languages that allow users to write explicit, format conceptualizations of domain models” [15]. It allows to semantically describe a system, and represent it into a model that could then be fed with data generated by the system. Entities can then be stored based on the ontology model which creates a database that could easily be accessed and read by humans thanks to its structure. In additive manufacturing there have been several eﬀorts working on ontology to represent the AM processes. There are diﬀerent ways to deﬁne the AM processes either trying to describe the entirety of the AM processes [16, 17] or either focusing a speciﬁc analysis of the AM processes in order to orient the model [18, 19]. Sanﬁlippo et al. focused on developing an ontology in order to represent the AM product lifecycle. They aim to describe all the AM processes that would result in the creation of the physical part [20]. Samyeon at al.

focused on developing an ontology in order to represent the AM product lifecycle. They aim to describe all the AM processes that would result in the creation of the physical part [20]. Samyeon at al. focused on developing an ontology that would focus on describing the data in order to support manufacturability analysis. This ontology is oriented in a way that it would be easier to determine the manufacturability of a part based on the data of the AM processes [21].Hierarchical data formats (HDF5) consists of a hierarchical model representing the system, a ﬁle format for storing and managing raw ﬁles generated from the system and meta-data describing the data contained in the raw ﬁles. It supports a variety of data types and is designed for managing huge volumes of data. In order to create HDF5 ﬁles the ﬁrst step is to think about the hierarchy in which the data will be segmented.

It supports a variety of data types and is designed for managing huge volumes of data. In order to create HDF5 ﬁles the ﬁrst step is to think about the hierarchy in which the data will be segmented. Then, once the hierarchy has been chosen the raw ﬁles can be stored in the corresponding leaf of the hierarchy [22]. This hierarchical data representation allows to create a tree structure that could eﬃciently segment huge dataset for easy analysis and handling [23]. The beneﬁt of using such format is the ability to eﬃciently store in an organized manner the data retrieved from the system and also manage the raw ﬁles generated.2. AM data challengesThe speciﬁcity of AM processes reside in the nature of the material being created. There is not one speciﬁc material that could be relied on during the entire lifecycle of the part.

There is not one speciﬁc material that could be relied on during the entire lifecycle of the part. On the contrary throughout the diﬀerent processes diﬀerent types of material are created which can be challenging to track and follow the evolution of such materials.Throughout the AM processes, a lot of diﬀerent data types and data formats are used and created. This results in challenges when trying to manage and analyze such huge and heterogeneous dataset. The AM product lifecycle is composed of diﬀerent stages. Each stage encompasses speciﬁc types of data and formats that have1552been developed in order to answer each AM process's needs [2]. The ﬁrst stage is the design and focuses on the part model, the recommendation, the topology, the design optimization, the tolerancing, the manufacturability assessment and the material selection. This stage aims to create a digital representation of the part and specify the requirement for the AM processes.

This stage aims to create a digital representation of the part and specify the requirement for the AM processes. The second stage is the process planning and focuses on predicting process response, property and performance response, and process optimization. This stage aims to ensure that the part could be able to be manufactured by preparing the manufacturing plan based on the part design and requirements from the previous stage. The third stage is the build and focuses on anomaly and defect detection, failure detection, machine condition monitoring, real time control, in-situ process monitoring and control. This stage is when the part is actually being manufactured and aims to monitor the manufacturing process. The fourth stage is the post process and focuses on post process control monitoring and control, process end point identiﬁcation, real time control. This stage aims to prepare and ﬁnish the part in order to meet the design requirements.

This stage aims to prepare and ﬁnish the part in order to meet the design requirements. The last stage is the testing and validation and focuses on surface metrology, defect detection and classiﬁcation. This stage allows the analysis of the resulting part and the comparison against the design requirements.The heterogeneity of data came from the huge diﬀerent types of data generated during the AM processes which makes it diﬃcult to eﬃciently handle, store and analyze. The geometry of the part can be described using the tessellated geometry, the point cloud geometry, the voxelized diﬀerent models : types generate speciﬁc data formats. The part material can also be geometry. Each of these diﬀerent characterized using diﬀerent the powder microstructure, the powder size distribution, the powder morphology, the powder rheology, the powder thermal properties, the powder optical absorption, the powder optical emissivity.

The AM process can be simulated through diﬀerent angles such as Powder simulation, melt pool simulation, build simulation. The in-situ monitoring generates multiple types of data from : layer wise staring imaging, acoustic monitoring, melt-pool monitoring, layer-wise monitoring. After the machining process, the ex-situ monitoring also generates multiple types of data based on : the microstructure, the X-ray topography.the solid geometry,types of data :the material chemistry,the powder chemistry,Thus, there is a need to deﬁne a structure that would enable the easy analysis and handling of such datasets and would also take into account the heterogeneity of data types generated throughout the diﬀerent AM processes.3. MotivationAM generates big heterogeneous dataset during its whole product lifecycle making it hard to manage and thus to analyze. The status quo for storing raw ﬁles through AM processes is either regular folders, or ﬁles format such as Excel.

The status quo for storing raw ﬁles through AM processes is either regular folders, or ﬁles format such as Excel. However, in order to apprehend the full potential of AM processes and enhance analytics opportunities, it is crucial to organize and store these big datasets in an eﬀective, usable way. The structures to register, capture and analyze the data are separated and not synchronized which result in time-consuming, labor-intensive data handling and data pre-processing. Moreover, ﬂat structures are used to describe and capture heterogeneous datasets which adds another diﬃculty for data curation.Thus, thanks to its hierarchical structure HDF5 could be a key enabler to represent AM data. Hierarchies can be created to store eﬃciently and comprehensively huge datasets allowing for easy access of AM data and then enhanced analysis.

Hierarchies can be created to store eﬃciently and comprehensively huge datasets allowing for easy access of AM data and then enhanced analysis. HDF5 also allows for chunking of data, thus based on the hierarchy, the dataset can be separated in diﬀerent chunks for faster and easier use.III.Proposed approach for AM hierarchical data organization1. Description of the approachOur goal for developing HDF5 data organization for AM is to be able to manage the diﬀerent datasets generated from the AM processes by coordinating and registering data to facilitate analysis. The structures for1553capturing and analyzing will be correlated using a more eﬃcient data management organization. The diﬀerent data types and formats used during the multiple processes would be reconciled and represented in a unique structure.

The diﬀerent data types and formats used during the multiple processes would be reconciled and represented in a unique structure. The data would be compose and compartment, for speciﬁc analysis, in order to ensure that only the relevant data is handled for analysis needs.Our approach will help to organize and store big datasets in an eﬀective way so that raw ﬁles are stored in hierarchies making it easy to access and curate. It will support data heterogeneity by enabling the storage of diﬀerent data types and formats in the same hierarchy. It will facilitate the handling of the data using chunks in order to separate the hierarchies in smaller datasets that are easier to analyze. Finally, it will facilitate data analysis by correlating the data registration structure and the data analysis structure.2. Hierarchy developmentIn order to develop a coherent hierarchy, some rules and guidelines must be deﬁned to ensure the validity and usability of the HDF5 ﬁle.

Hierarchy developmentIn order to develop a coherent hierarchy, some rules and guidelines must be deﬁned to ensure the validity and usability of the HDF5 ﬁle. First, the hierarchies should be developed in a way that all the leaves of the hierarchy should lead to heavy data. Thus, all the buckets containing the data in the leaves of the hierarchies should contain multiple data or datasets. The taxonomy describing the hierarchy should be intuitive in order to facilitate the curation and the chunking of the data. It should ensure that the taxonomy is clear enough to allow for easy identiﬁcation of the underlying datasets and to prevent any errors of interpretation. The taxonomy should also be deliberate and developed in order to answer AM data challenges. The description of the hierarchy following an analysis purpose will allow to bridge the data registering and the data analysis structure, thus helping the data curation and data processing.

The description of the hierarchy following an analysis purpose will allow to bridge the data registering and the data analysis structure, thus helping the data curation and data processing. The hierarchy should also be agnostic of the process technology. The AM technologies characteristics should not be of importance for the taxonomy development. This will allow to create more generic hierarchies that could be reusable and understandable. Finally, the hierarchy should be developed from a global point of view. It should not be speciﬁc to a use case or an application.Once the taxonomy rules have been identiﬁed, the next focus will be on identifying the data to represent. Depending on the AM data challenge’s needs, the relevant data should be selected. Then, based on this dataset, diﬀerent data types and formats, from multiple AM processes, will be involved.

Depending on the AM data challenge’s needs, the relevant data should be selected. Then, based on this dataset, diﬀerent data types and formats, from multiple AM processes, will be involved. Finally, having the datasets and information about the data types and formats, the last step is to think of the composition of the data and how to facilitate the handling of such dataset.The last step of the hierarchy development is the development of the taxonomy. Once, clear rules for the development and the data have been identiﬁed, the taxonomy can be deﬁned. The development of the taxonomy should be driven by context and perspective speciﬁc AM data challenges. The focus of the hierarchy should be chosen in order to either develop a goal oriented or generalized hierarchy. Finally, the hierarchy should answer data management needs and help correlate the data analysis and data storage representation of AM data.3.

Finally, the hierarchy should answer data management needs and help correlate the data analysis and data storage representation of AM data.3. Hierarchies examplesIn order to illustrate our approach, we have worked on four diﬀerent examples of hierarchies that would allow to answer AM data challenges. The ﬁrst hierarchy that we are focusing on is the data curation aspect. This hierarchy aims to facilitate the curation of the data by making it easy to store, easy to access and easy to understand. This perspective is important since there are so many diﬀerent types of format and data produced from and during AM processes. The goal would be for the user to instinctively access and store the data. Thus, the hierarchy needs to be optimized in order that there is no confusion on where to store the data and no confusion on where to retrieve the data. The leaves of the hierarchy should be speciﬁc enough so that the type of data in it can be stored only here and nowhere else.

The leaves of the hierarchy should be speciﬁc enough so that the type of data in it can be stored only here and nowhere else. However, it doesn’t mean that only one type of data can be stored on a speciﬁc leaf but multiple data types can be in it as long as they can’t be anywhere else.1554Figure 2 : Data curation hierarchyThis hierarchy aims to describe the AM processes and the data generated in order to create a dictionary of data or a data catalog that could describe and contain all the AM data. The hierarchy is separated into diﬀerent type of information, design, sensor, build, process, machine, material. Each category contains unique information of the processes or part at this speciﬁc point of time during the AM product lifecycle. This hierarchy is a data-centric approach that aims to describe the data in AM processes. However, the hierarchy is not based on data formats, e.g.

This hierarchy is a data-centric approach that aims to describe the data in AM processes. However, the hierarchy is not based on data formats, e.g. Image data, Signal data, Geometry data, since it can be diﬃcult to access from an AM perspective and it does not represent AM processes. This hierarchy would allow for easy publication and exchange of the dataset and with a common hierarchy describing the data it would be easier to work in collaboration. The hierarchy would be self-describing allowing for easy use and maintenance by third party. This would also serve as a data catalog to eﬃciently store the data. This would allow for easy storage and to be a base for the data mining in order to create more advanced data analysis.

This would also serve as a data catalog to eﬃciently store the data. This would allow for easy storage and to be a base for the data mining in order to create more advanced data analysis. It would allow for an easy overview of the data available and would ensure that the whole scope of the AM processes could be apprehended.The second hierarchy focuses on the comparison between the virtual part and the physical part, and on the build replication in production scenarios. The goal is to compare the virtual and physical, observed parts and processes. The main challenge of this hierarchy is to analyze what we are comparing against and what we are validating.Figure 3 : Part validation hierarchyThe hierarchy of this perspective will create 2 segments containing respectively the virtual part and the physical part. Each side of the hierarchy will focus on the data associated with either the virtual or the physical part.

Each side of the hierarchy will focus on the data associated with either the virtual or the physical part. A clear distinction between the 2 parts must be made throughout the processes. Each AM process is analyzed in order to extract the virtual data from the physical data. The data types are not in the center of the hierarchy but each data type is stored in a way that would facilitate the validation of each process. This hierarchy is suited to compare the expected design to the actual part. It allows part validation and comparison against design requirements. Having the data segmented in between expected and measured would allow for easy identiﬁcation of discrepancies and issues in the resulting part. This hierarchy would also allow for process analysis.

Having the data segmented in between expected and measured would allow for easy identiﬁcation of discrepancies and issues in the resulting part. This hierarchy would also allow for process analysis. Since the requirements for each process are deﬁned, it would be easy to analyze how the process has been made and identify potential errors or optimization capabilities.The third hierarchy would aim to follow the evolution of the part through the AM Product Lifecycle. Thegoal would be to track the changes and evolution of the part. e.g. spatial volumes, discrete spaces…, over time.1555Figure 4 : Product Lifecycle hierarchyThis hierarchy would be separated by each stage of the AM Product Lifecycle. The data generated in each stage would be stored in the corresponding part of the hierarchy allowing to have in the same place the data related to a speciﬁc AM process.

The data generated in each stage would be stored in the corresponding part of the hierarchy allowing to have in the same place the data related to a speciﬁc AM process. This hierarchy doesn’t aim to capture the “witness specimen” but trace the evolution of the part and what the part looks like at any given time. It is a time-centric approach that captures the state of the part at any given time. It follows as the part matures throughout its lifecycle. This hierarchy would allow the same type of data being stored in diﬀerent parts of the hierarchy but where the data will be stored depends on which processes the type of data is associated with. This hierarchy would allow to follow the manufacturing of the part and monitor the manufacturing process. Having the hierarchy separated based on the lifecycle stage would allow for users to follow the part even if some part of the processes is externalized.

Having the hierarchy separated based on the lifecycle stage would allow for users to follow the part even if some part of the processes is externalized. The data from contractors could be stored in the corresponding stage and user could verify the state of the part in its lifecycle. Since each data of the AM product lifecycle is stored based on its stage, this would allow to trace and better control the system. All the data related to the same process will be stored in the same place allowing for easy analysis of a speciﬁc process.This perspective aims to standardize the creation of the material based on the build. The goal is to be suited to be queried by material database. I will develop material information with the design and allow to follow the material evolution throughout the AM processes.Figure 5 : Material evolution hierarchyThis hierarchy will follow the AMMD model and aims to record and follow the evolution of the material throughout the product lifecycle.

The hierarchy will be material-centric and segment the AM processes such as for each part of the hierarchy the material at a speciﬁc point of time will be described and the data related to this material will be stored together. The diﬀerent stages of the material will be segmented and the corresponding processes that allowed the creation of such material will be described in each part.1556This hierarchy allows for analysis of the evolution of the material. Thus, this hierarchy would allow for control of the material quality and identiﬁcation of defects in the material. Having the diﬀerent stages of the evolution segmented will also allow to identify where the defect might happen and how to improve or modify the process in order to prevent such defect.Case study applied to Digital Twin The proposed approach to create a HDF5 hierarchy in order to answer AM data challenges could be applied and used as a basis to build Digital Twin applications.

The corresponding hierarchy would contain the information needed to realize the Digital Twin and analysis could be directly realized using the curated and pre-processed data stored in HDF5.IV.1. Description of the case studyWe showcased our approach with a case study obtained from a LPBF build in the Additive Manufacturing Metrology Testbed (AMMT) at NIST [24]. The dataset includes the tessellated geometry of the part, the in-situ monitoring and ex-situ monitoring data. The tessellated geometry contains the 3d representation of the part and the 2d layer-wise geometry. The in-situ monitoring data contains the melt-pool images, the layer-wise melt pool and layer-wise surface.

The in-situ monitoring data contains the melt-pool images, the layer-wise melt pool and layer-wise surface. The ex-situ monitoring data contains the XCT images and as-manufactured surface and edges.Part geometryIn-situ monitoringEx-situ monitoringMelt pool imagesXCT imagesTessellated geometryLayer-wise 2d geometryLayer-wise melt poolLayer-wise surfaceAs-manufactured surface and edgesFigure 6 : Case study dataset representation [25].2. Hierarchy for Digital TwinFor this case study, the focus is on analyzing the resulting as-manufactured part and identifying potential defects such as pore structures inside the layer of the part. Thus, the previous proposed hierarchy for part validation (see Figure 3) could be adapted in order to answer the needs to qualify the part.

Thus, the previous proposed hierarchy for part validation (see Figure 3) could be adapted in order to answer the needs to qualify the part. Since the layer-wise geometry and monitoring data will be analyzed, the geometry must be reﬂected in order to enable easy linking of 2d to 3d data.Figure 7 represents the part validation hierarchy adapted in order to store the datasets generated by the case study experiment. Each dataset is stored in a leaf of the hierarchy that represents the data contained in the corresponding dataset.1557Figure 7 : Case study hierarchy for part validation and Digital Twin development3. Hierarchy applicationOnce the dataset has been properly described and stored in the HDF5 hierarchy, multiple applicationscan be done in order to build replicas and Digital Twins of the part. Since the data is already pre-processed and linked based on layer-wise and geometry data, the analysis on the dataset would be facilitated.

Since the data is already pre-processed and linked based on layer-wise and geometry data, the analysis on the dataset would be facilitated. For example, using the XCT layer-wise images, Figure 8 shows the 3d XCT volume can be reconstructed which would render data features for the as-manufactured part that could be compared against the tessellated geometry. Defects can be identiﬁed inside the diﬀerent layers of the part and pore structures can be analyzed. The resulting defects that have been found can be compared against the corresponding layer-wise melt-pool and surface to identify potential areas that would lead to such defects.

The resulting defects that have been found can be compared against the corresponding layer-wise melt-pool and surface to identify potential areas that would lead to such defects. Based on these areas, decisions can be rendered, such as geometry optimization, scan strategy modiﬁcations, parameters adjustments…, in order to prevent the observed and identiﬁed defects.Figure 8 : 3d XCT model reconstructed with defects detectionV.Conclusion Data generated through AM processes lead to the creation of huge dataset that are heterogeneous intypes and formats. Thus, big data management is becoming a growing challenge in AM that impedes the use of the full potential of data analysis. Diﬀerent methods have been developed in order to help the handling and curation of AM dataset but there is still a lack of common structures between the data registration and the data analysis.

Diﬀerent methods have been developed in order to help the handling and curation of AM dataset but there is still a lack of common structures between the data registration and the data analysis. In this paper, we have focused on developing a new approach that uses HDF5 in order to store the registered heterogeneous data from diﬀerent AM processes and allow to directly apply data analysis using the1558same structure. We have showcased four diﬀerent hierarchies that answer speciﬁc AM data challenges and would help easy handling and partitioning of big datasets. Finally, we have demonstrated our approach with a Digital Twin application that would allow to use HDF5 as a holder for the data to represent the Digital Twin of the part.

Finally, we have demonstrated our approach with a Digital Twin application that would allow to use HDF5 as a holder for the data to represent the Digital Twin of the part. Example applications such as defects detection and design optimization can then be easily applied thanks to the coherent hierarchy structure.In future work, we planned on developing a full data registration structure that would contain largerdataset and showcase the potential of using HDF5 for automated data analysis. The goal will be to focus on the metadata of the datasets in order to describe the data and generate comprehensive pipeline to facilitate data analysis.References :[1] Gibson, Ian, David W. Rosen, Brent Stucker, Mahyar Khorasani, David Rosen, Brent Stucker, and Mahyar Khorasani. Additive manufacturing technologies. Vol. 17. Cham, Switzerland: Springer, 2021. [2] Razvi, Sayyeda Saadia, Shaw Feng, Anantha Narayanan, Yung-Tsun Tina Lee, and Paul Witherell.

at Transfer Analysis of a Gas-Jet Laser Chemical Vapor Deposition (LCVD) ProcessChad E. Duty, Brian T. Fuhrman, Daniel L. Jean, and W. Jack Lackey Rapid Prototyping and Manufacturing Institute Woodruff School of Mechanical Engineering Georgia Institute of Technology Atlanta, GA 30332AbstractThis paper describes the development of a computer model used to characterize the heat transfer properties of a gas-jet LCVD process. A commercial software package was used to combine heat transfer finite element analysis with the capabilities of computational fluid dynamic software (CFDS). Such a model is able to account for both conduction and forced convection modes of heat transfer. The maximum substrate temperature was studied as a function of laser power and gas-jet velocity.IntroductionLaser chemical vapor deposition (LCVD) is a relatively new manufacturing process that holds great potential for the production of small and complex metallic and ceramic parts.

A sophisticated gas-jet LCVD system has been designed and developed at Georgia Tech. The system promises increased geometric flexibility and deposition rates and is designed to function as a rapid prototyping system. The system is currently under development to understand and optimize key processing parameters. Since CVD is a thermally activated process, the most important process variable is temperature. In a laser-heated process such as pyrolytic LCVD, the temperature field is restricted to a micron scale and can vary by an order of magnitude over the diameter of the laser spot. Deposition rates typically follow an Arrhenius relationship that is exponential with respect to temperature, so it is critical to document and understand these two- dimensional temperature variations.BackgroundA number of theoretical and experimental studies have attempted to model and understand the role of temperature in the LCVD process.

Before the advent of LCVD, Kokorowski et al.1 predicted temperature distributions in a substrate during continuous wave laser annealing. Conduction was the only mode of heat transfer considered, but he did account for temperature dependent material properties. Many studies2-4 followed in the next 50 years, but few of the models allowed for forced flow convection and many did not account for temperature- dependent material properties.461Steen's5 numerical model, developed in 1977, demonstrated that convection and radiant surface losses were significant in a temperature sensitive process. He introduced convection losses into the model by including a non-reacting gas jet that was coaxial with the laser beam. By restricting the jet to be coaxial, he preserved the axial symmetry of his model, making analytical calculations possible.

By restricting the jet to be coaxial, he preserved the axial symmetry of his model, making analytical calculations possible. Steen studied the thermal history of the substrate surface with respect to various process parameters, including the jet Reynolds number, the distance between the substrate and the nozzle mouth, the jet temperature, and the surface reflectivity.In 1989, Mazumder and Kar6 introduced a three-dimensional model that incorporated all three mechanisms of heat transfer (conduction, convection, and radiation). The analysis employed a technique of uniformly moving finite slabs to synthesize a moving laser beam at a constant velocity.

The analysis employed a technique of uniformly moving finite slabs to synthesize a moving laser beam at a constant velocity. In contrast to the previous infinite and semi-infinite geometric models, this analysis was much more applicable to smaller substrates whose overall dimensions were on the order of the laser spot diameter.In 1998, Yu and Duncan7 proposed a three-dimensional finite element model that was able to account for the natural convection fluid flow that results from laser heating. The model included conduction, convection, and radiation heat transfer mechanisms. A buoyancy-driven flow was determined by the Boussinesq principle and demonstrated during the horizontal growth of a fiber. Yu and Duncan assumed incompressible fluid, laminar flow, and steady state conditions.

Yu and Duncan assumed incompressible fluid, laminar flow, and steady state conditions. The effects of the natural convection were demonstrated in terms of temperature gradients and associated growth rates.It is apparent that a significant amount of work has been performed in the area of LCVD process and thermal modeling. The analytical models and numerical analyses have progressed considerably over the years, however none of the models allow for a forced directional gas flow.Model DevelopmentPhysical System DescriptionGeorgia Tech's gas-jet LCVD system is designed to optimize each aspect of the LCVDprocess with respect to build time, accuracy, resolution, and flexibility. The system has been described in detail in previous papers8,9 and is shown schematically in Figure 1. A small reaction chamber houses a substrate on top of a moving and rotating shaft.

The system has been described in detail in previous papers8,9 and is shown schematically in Figure 1. A small reaction chamber houses a substrate on top of a moving and rotating shaft. High resolution, three- dimensional patterns can be created by numerically controlling the movement of the rotating stage assembly. By travelling in a spiral pattern, fast scanning speeds can be achieved without the high reversal forces typical of raster scanning. Reagent gases are introduced into the reaction chamber via a small diameter gas jet that impinges on the substrate at a 45° angle. The high- velocity jet generates a thin boundary layer for the diffusion of reagent gases, theoretically enhancing the chemical kinetics of the reaction and increasing the deposition rate. A carbon dioxide laser heats a spot on the substrate above the threshold temperature required to initiate pyrolytic LCVD reactions.

A carbon dioxide laser heats a spot on the substrate above the threshold temperature required to initiate pyrolytic LCVD reactions. The laser has a maximum power rating of 100 W and is focused to a beam diameter of 200 µm.462CO2 Laser BeamGas JetUpper Reaction ChamberDepositSubstrateVacuum LineFigure 1. Cross Section View of LCVD Reaction ChamberModel Geometry and Boundary ConditionsThe reaction chamber geometry has been simplified in the finite element model in order to make calculations more efficient. Since the model is symmetric about the axis of the reagent jet, only one half of the area surrounding the gas-jet and laser spot is represented. The geometry and boundary conditions of the wedge-shaped model are illustrated in Figure 2. The substrate consists of a _ inch thick piece of graphite. A thin-walled tube geometry is suspended in the argon atmosphere above the substrate to represent the reagent gas-jet.

The substrate consists of a _ inch thick piece of graphite. A thin-walled tube geometry is suspended in the argon atmosphere above the substrate to represent the reagent gas-jet. The end of the tube is separated from the center of the laser spot by 3/16 of an inch. The laser spot is represented as aVel = 50 m/s Temp = 300 KATemp = 300 KArgon AtmospherePress = 1 atmI.D. = 1 mmSee Cutout A Gas-JetHeat Power = 10 WBSymmetrySee Cutout BCarbon SubstrateTemp = 600 KLaser Spot Dia = 200 µmFigure 2. Model Dimensions (cm) and Boundary Conditions463small cylinder embedded in the carbon substrate. The cylinder protrudes from the substrate by 20 nm, simulating the first stages of deposit growth. This also allows for a complementary face within the argon atmosphere to be used to specify boundary conditions.The physical properties of the laser cylinder and the carbon substrate are specified as grade ATJ isomolded graphite.

The properties of the graphite and the argon are modeled as constant with respect to temperature. The fluid flow properties of the model are specified at a single inlet and outlet. Argon enters the tube with a specified velocity and temperature (see cutout 2-A). A typical inlet velocity used in practice is 50 m/s. The outlet is represented by a total pressure boundary of 1 atmosphere on the back face of the argon atmosphere. The laser power is represented as heat power boundary condition across the top face of the laser cylinder (see cutout 2-B). Since the heat power is applied to only one-half of the laser spot, the effective laser power is twice this amount (i.e. 10 W of applied heat power translates to a laser power of 20 W, assuming that the substrate is a blackbody). Physically, the laser in the LCVD process has a Gaussian distribution, but it is restricted in this model to a flat distribution.

Physically, the laser in the LCVD process has a Gaussian distribution, but it is restricted in this model to a flat distribution. Constant temperature conditions were also set on the top and bottom surfaces of the model to respectively simulate room temperature (300 K) and the resistive heating element (600 K).ABFigure 3 shows the result of the automatic meshing routine used by Cosmos/Works®. Mesh control conditions were set on the symmetry face of the laser cylinder and the terminating face of the gas-jet tube. The grid spacing on the laser face was restricted to 20 µm with a geometric growth rate of 1.15 over 25 layers. The grid spacing on the tube face was restricted to 100 µm with a geometric growth rate of 1.15 over 25 layers. These conditions resulted in a very fine mesh in the regions of interest while limiting the overall model to just under 16,000 nodes.Figure 3.

These conditions resulted in a very fine mesh in the regions of interest while limiting the overall model to just under 16,000 nodes.Figure 3. Mesh spacing on symmetry face for (a) tube and laser and (b) close-up of laser cylinder.Model Evaluation and ResultsThe model was evaluated using a commercial code, Cosmos/Flow®, that is designed to solve the mathematical equations that represent heat transfer and momentum transfer in a moving fluid. The current model represents an initial phase in the development of a more sophisticated thermal model of the LVCD process. In this preliminary phase, conduction and forced convection will be the only modes of heat transfer considered. The effects of radiation and natural convection (buoyancy-driven flow) will be ignored. The internal flow is considered turbulent and fully compressible.464Figure 4 illustrates the flow velocity along the symmetry plane for an inlet velocity of 50 m/s.

The internal flow is considered turbulent and fully compressible.464Figure 4 illustrates the flow velocity along the symmetry plane for an inlet velocity of 50 m/s. The laser spot is visible within the substrate and there appears to be significant flow just above the surface of the laser spot. Figure 5 shows how the temperature in this plane is affected by the gas-jet with no effects from the laser beam. The dominant mode of heat transfer around the laser spot is apparently conduction, since the surface temperature of the substrate is equal to the heater temperature of 600 K. Figure 6 is a contour plot of the gas flow just above the surface of the substrate (separation distance of ~100 µm). The location of the laser spot is shown as a white circle. Figure 7 demonstrates how the temperature within the laser spot varies when the laser power is 20 W and there is no argon flow in the chamber. The maximum temperature in this case is about 1700 K.

Figure 7 demonstrates how the temperature within the laser spot varies when the laser power is 20 W and there is no argon flow in the chamber. The maximum temperature in this case is about 1700 K. Since the isotherms are not concentric circles, the symmetry conditions of the model may not be appropriately defined.AABSubstrateBSubstrateFigure 4. Flow Contours of Symmetry Plane. Inlet Velocity = 50 m/s. VelA = 35 m/s VelB = 0 m/sFigure 5. Temperature Contours of Symmetry Plane. Power = 0 W, Vel = 50 m/s. TA = 300 K TB = 600 KABFigure 6. Flow Contours above Laser Spot. Inlet Velocity = 50 m/s. Center Vel = 10 m/s Edge Vel = 0 m/sFigure 7. Temperature Contours of Laser Spot. Power = 20 W, Vel = 0 m/s. TA = 900 K TB = 1680 K465Figures 8 through 10 below show temperature variations across the laser spot at several values for the inlet flow velocity. In each case, fluid flow is from the right and the laser power is constant at 20 W.

In each case, fluid flow is from the right and the laser power is constant at 20 W. The resulting plots are nearly identical, both in terms of absolute temperatures and distributions. Surprisingly, this suggests that convection heat transfer plays an insignificant role in the LCVD process. Another interesting feature is the existence of two distinct hot pockets along the centerline of each spot. The cause of these is unclear, though it can most likely be attributed to ill-defined symmetry conditions of the model.AABFigure 8. Temperature Contours of Laser Spot. Power = 20 W, Vel = 30 m/s. TA = 900 K TB = 1700 KFigure 9. Temperature Contours of Laser Spot. Power = 20 W, Vel = 50 m/s. TA = 900 K TB = 1700 KAABFigure 10. Temperature Contours of Laser Spot. Power = 20 W, Vel = 100 m/s. TA = 900 K TB = 1700 KFigure 11. Temperature Contours of Laser Spot. Power = 80 W, Vel = 50 m/s.

Temperature Contours of Laser Spot. Power = 20 W, Vel = 100 m/s. TA = 900 K TB = 1700 KFigure 11. Temperature Contours of Laser Spot. Power = 80 W, Vel = 50 m/s. TA = 2000 K TB = 5000 K466BBFigure 11 on the previous page illustrates typical temperature variations at the higher absolute the relative distribution across the laser spot is nearly identical to those with lower laser powers. Figure 12 plots the maximum temperature within the laser spot for a range of laser powers. The result of elevating the substrate temperature by increasing is hardly unexpected. The surprising outcome is the linear correlation of the data -- it is perfectly linear with an intercept of 600 K and a slope of 56 K/W.laser powers.Although temperatures are higher,thelaser powerDiscussionThe results suggest that the gas flow velocity has a negligible effect on the absolute temperature and temperature distribution within the laser spot.

Therefore, it appears that convection is not a significant mode of heat transfer under the conditions of this model. This result is unexpected considering the proximity of a high velocity gas-jet. However, a quick analysis shows that the energy absorbed by the passing argon gas per unit time is about 1% of the power supplied by the laser.Consider that argon is flowing out of the nozzle at 50 m/s. Given the diameter of the nozzle and a density for argon of 0.5 kg/m3, the mass flow out of the gas-jet is 4E-05 kg/s. Suppose that as the jet expands and approaches the 200 µm laser spot, about 1% of the gas passes directly over that area. From the results of the above model (Figure 9), we see that the maximum temperature across the laser spot is about 1700 K for a laser power of 20 W.

From the results of the above model (Figure 9), we see that the maximum temperature across the laser spot is about 1700 K for a laser power of 20 W. Given that the heat capacity of argon is 520 J / kg / K and the original substrate temperature is 600 K, it is possible to calculate the energy absorbed by the argon gas as it passes over the laser spot.P = D T (dm/dt) cp = (1700 - 600 K) (4E-07 kg/m3) (520 J/kg/K) = 0.23 J/s = 1.1 % of 20 WThis is a rough estimate, but it supports the idea that convection is not a dominant mode of heat transfer under these conditions. The linear behavior of Figure 12 can be interpreted more appropriately once conduction is identified as the sole source of heat transfer (remember that radiation is not included in this model). Conservation of energy requires that the heat flux generated by the laser must equal the heat flux dissipated by conduction. This is expressed below along with the basic form of the conduction equation.

This is expressed below along with the basic form of the conduction equation. The terms for conductivity and area remain constant with respect to temperature in this model. Therefore, the conductivity equation can be rearranged into the linear y = m x + b format as shown below. With this perspective, the linear relationship of Figure 12 is not so surprising -- it merely points to a conduction-dominated process. Furthermore, the y-intercept equals the original substrate temperature, 600 K.467fluxlaser = fluxconduct = k / A ( D T ) Ł Tmax = (A / k) flux + ToConclusionsLCVD is a process dominated by temperature. It is imperative to study and accurately model the temperature variations in and around the laser spot. The current model provides a rough estimate and serves as an important first step in the development of a more accurate model. A surprising conclusion from this model is the relatively small impact of convection on the heat transfer of the gas-jet LCVD process.

A surprising conclusion from this model is the relatively small impact of convection on the heat transfer of the gas-jet LCVD process. We will attempt to verify this result through future versions of the model that will include temperature-dependant material properties, natural convection, radiation, and a Gaussian beam distribution. One should hesitate, however, to discount convection as a possible source of heat transfer based on this model alone.AcknowledgementsThis work was supported by the Engineering Research Program of the Office of BasicEnergy Sciences at the U. S. Department of Energy and the Georgia Institute of Technology.References1. S.A. Kokorowski, G.L. Olson, and L.D. Hess. Proceedings of Laser and Electron-Beam Solid Interactions and Material Processing. J.F. Gibbons, L.D. Hess, and T.W. Sigmon, eds. New York 1955. pp. 139-146.2. J.E. Moody and R.H. Hendel. “Temperature Profiles Induced by a Scanning cw Laser Beam.” Journal of Applied Physics 53. 1982. p.

at Transfer Manipulation for Precision Droplet Manufacturing: Simulation and ExperimentMatthew Michaelis*, Jun Zhu*, and Melissa Orme+Graduate student researcher, +Professor Dept. of Mech. and Aerospace Engineering University of California, Irvine Irvine, California 92697-33975Abstract The solid free-form fabrication (SFF) of arbitrary metallic components is accomplished with controlled deposition of nano-liter molten droplets generated by capillary stream break-up and deposited at rates on the order of 10,000 drops/second. By varying the droplet arrival temperature, deposition rate, and substrate temperature in both the simulation and experiment, we seek to obtain a thorough understanding of the heat transfer phenomena that governs the SFF component quality.

Of specific interest is the removal of inter-splat boundaries in order to achieve a high quality component, characterized by a uniform and fine microstructure, by having newly arriving drops remelt a thin layer of the previously deposited and solidified material. A numerical model, which simulates the heat transfer manipulation, is used to understand and guide the process development.Introduction The ability to build arbitrary 3D objects directly from CAD data without the need for molds, a process often referred to as rapid prototyping (RP), has been around for a number of years1. Precision droplet manufacturing (PDM) is a molten metal droplet based RP technique being developed at UCI that differs from many other metal RP technologies. Specifically, it is capable of producing near net-shape metal components that exhibit high accuracy, improved material properties, fast build times, and virtually no warpage.

Specifically, it is capable of producing near net-shape metal components that exhibit high accuracy, improved material properties, fast build times, and virtually no warpage. Additionally, since PDM produces fully dense components, no secondary operations such as post firing or infiltration are required.The PDM process utilizes a capillary stream of molten metal that is broken up into highly uniform drops, which can be individually manipulated via electrostatic deflection on a drop-to-drop basis . Electrostatic droplet deflection in combination with xyz substrate motion is used to build arbitrary 3D objects.2 Components can be grown at rates up to 40 mm3 per second since the 200 µm diameter droplets are generated at a rate on the order of 10,000 droplets per second. At this rate, a solid 1.0 cm3 cube can be produced in 24 seconds. Accuracy and surface roughness are determined by the stream stability and splat characteristics of the drops.

At this rate, a solid 1.0 cm3 cube can be produced in 24 seconds. Accuracy and surface roughness are determined by the stream stability and splat characteristics of the drops. Figures 1 and 2 are photographs of components that were processed with PDM.Figure 1 Thin wall clover built with PDM368Having previously deposited drops solidify before the next drop arrives is vital so that geometric integrity is maintained. However, enough thermal energy must be delivered with each new drop to remelt a thin layer of the previously deposited material to help remove inter-splat boundaries and thus create a strong inter-splat bond. Remelting promotes inter-droplet adhesion by allowing mixing and therefore should minimize delamination.

Remelting promotes inter-droplet adhesion by allowing mixing and therefore should minimize delamination. Therefore, superheated drops in combination with a cool substrate will allow remelting yet pull heat away fast enough that the splat can solidify before the next splat arrives.To to achieve quality components, a numerical simulation is developed and compared with experiments. With these two goals in mind, the model can be used to predict optimum parameters that yield the maximum deposition build height in relation to droplet generation frequency, droplet arrival temperature, and substrate temperature.investigatethe conditionsrequiredFigure 2 Braids built with PDMNumerical Model A 1D model is used to simulate the bulk material deposition and heat diffusion process associated with PDM. Heat transfer and solidification are simulated as droplets are added at regular intervals. The first drop lands on a finite substrate and subsequent drops stack on the previous drop.

Heat transfer and solidification are simulated as droplets are added at regular intervals. The first drop lands on a finite substrate and subsequent drops stack on the previous drop. Two moving boundaries are associated with PDM: (1) liquid/solid interface that varies continuously with time and (2) the upper adiabatic boundary that varies discretely as each new drop is added. The model allows re-melting of previous layers, but not the substrate. Thus, the liquid/solid interface is free to move in two directions. A constant is temperature boundary condition applied to the bottom of the substrate and an adiabatic condition is applied to the free droplet surface. A fixed numerical grid scheme is used.

A fixed numerical grid scheme is used. The basis for this formulation can be found in detail elsewhere.3stagnation pressure regulationdroplet generation waveform (shown for conventional sinusoidal disturbance)reservoirλλmolten metalvibrating plungercharge waveform (shown for charging every other ball)orificecontinuous stream of droplets is generated with an average droplet spacing of λcharge electrodeλ λdeflection electrodesF=qEcatcherCrank-NicolsonsubstrateThe adiabatic boundary condition atFigure 3 General schematic of PDM Apparatus illustrating the droplet generation, charge and deflection mechanisms.369the top surface is valid since the dominating heat transfer mode is conduction through the substrate.8 Additionally, it is assumed that the droplet deformation will not influence the heat diffusion process. This assumption is reasonable because of the high impact velocities associated with PDM.

This assumption is reasonable because of the high impact velocities associated with PDM. Besides, modeling the deformation would make the model unwieldy when analyzing hundreds or thousands of droplets (which is very realistic for PDM).9Experimental Apparatus Molten metal is jetted through a small orifice (100 µm) at the base of the liquid metal reservoir as shown in figure 3. This stream is broken into uniform molten metal drops through capillary stream breakup induced with a periodic pressure disturbance to the stream. The entire stream can be given an electrical charge with the charge electrode. When a drop breaks free from the stream, it retains whatever charge was on the stream at the time of separation. Thus, arbitrary charges can be applied to each drop by applying a varying voltage to the charge electrode that is synchronized with the droplet breakup perturbation. The arbitrarily charged drops then pass through deflection plates that generate a constant electric field.

The arbitrarily charged drops then pass through deflection plates that generate a constant electric field. Each drop will then be laterally deflected a distance proportional to its charge.Droplet stream phase and deflection calibration To help guarantee deposition control, it is important that charging pulses be in phase with the droplet generation waveform and that the droplet breakup location is relatively invariant. If neither of these conditions is satisfied, drops may only receive a fraction of their nominal charge value or even no charge at all, resulting in large targeting errors. Therefore, it is desirable to have a visual method so that noncompliance can be easily identified and eliminated. An innovative method has been developed that achieves this goal.A video camera that is oriented perpendicular to the deflection direction is used to observe the droplet stream.

An innovative method has been developed that achieves this goal.A video camera that is oriented perpendicular to the deflection direction is used to observe the droplet stream. Verification of the droplet formation location invariance, or stability, can be easily accomplished by applying a constant charge voltage and observing if all the drops are deflected the same amount. To verify synchronization between the charging and droplet generation waveforms, a charging waveform that is highly sensitive to phase shifts is used. To this end, a droplet charging waveform with a steep voltage gradient that is centered about zero volts is applied. Thus, when generation and charging are slightly out of phase the trajectory of neighboring drops diverge noticeably. Hence, neighboring drops are aligned only when the generation and charging waveforms are in phase.

Hence, neighboring drops are aligned only when the generation and charging waveforms are in phase. This method provides a true and readily observable indicator of phase relationship since it examines the final result of the charge input.Both calibration schemes can be implemented simultaneously by using a clipped sine wave for the charge waveform with a frequency that is 1/4 of the droplet generation frequency. The top of figure 4 shows how the charging waveform relates to the droplet stream when they are in phase. Note that every other drop has zero volts applied to it and is in a steep voltage gradient region as desired. In addition, every other drop is on the flat peak of the clipped sine wave and can be used to observe deflection stability. Since the zero voltage drops are in a steep voltage gradient region, a small phase shift will result in a large change in the charge voltage, resulting in large errors in droplet trajectory and placement as shown in the bottom of figure 4.

It is clear that the phase shift causes the center drops to no longer travel in tandem since they were given non-zero charges at the time of droplet generation. The calibration method involves “tuning” the phase370In PhaseDrop Generation WaveformDrop StreamCharge WaveformDeflected Drop StreamOut of PhaseDrop Generation WaveformDrop StreamCharge WaveformDeflected Drop StreamFigure 4 Calibration Schemedifference between the droplet generation waveform and charge waveform until the center droplets are observed to travel in tandem.Experiment For the current experiment, solder (60Sn-40Pb)7 droplets splat and form a cylindrical bead 450 µm thick on a 3.2mm thick aluminum substrate mounted on a rotating platform. The substrate rotates at a frequency of 1.3 Hz for 15 revolutions to build each component. The bottom surface of the substrate is held at a constant temperature. As the platform rotates, a bead of molten metal is deposited.

The bottom surface of the substrate is held at a constant temperature. As the platform rotates, a bead of molten metal is deposited. After one revolution, material is deposited on top of the previously deposited material. In this way a slinky shaped component as shown in figure 5 is built. The bonding between the “rings” of the component is of primary interest.To evaluate the inter-layer bonding a metallographic study of the specimens is conducted. Droplet arrival temperature is held constant while substrate temperature is varied. The grain structures of a sample is shown in figure 6 (discussed in more detail later). The bottom layers,371which were closest to the substrate, exhibit a fine grain structure that then transitions quickly to larger grain structure further away from the substrate. Towards the top the component looses geometric integrity.Results and Discussion Larger grains indicate that the solder is cooling slower than for smaller grains.

Towards the top the component looses geometric integrity.Results and Discussion Larger grains indicate that the solder is cooling slower than for smaller grains. This occurs since the substrate is being continuously heat by the arriving molten solder. Longer cooling times give a layer more of a chance to bond with the lower layer and a stronger bond results. It is theorized that the transition region show in figure 6 is where strong inter-layer bonding starts to occur. At some point, solidification slows such that a layer is unable to fully solidify before the next layer arrives and geometric integrity is compromised. All of these regions are illustrated in figure 6.Figure 7 shows the temperature history for the simulation through a succession of 15 layers, where the solid line is the substrate temperature profile and the dashed line is the temperature profile of the top surface of the last droplet deposited.

To compare the simulation with the experiment, the last layer to fully solidify before the next layer arrives is examined for the temperature range of interest. This point is indicated in figure 7. This point is chosen since, as shown in the micrograph of figure 6, the point at which the layers are no longer distinct is easily located. The simulation predicts when this point will occur for a certain substrate temperature as shown in figure 7. It can be seen from figure 8, which is a plot comparing the experimental results of the metallographic samples with the predictions, that the simulation is in good agreement with experiment.Experimental results are presented in figure 8 along with our simulation predictions. Even though there is some variation in the metallographic results the specific downward trend is clearly evident. The slope of the lines is the same within the experimental variations.

Even though there is some variation in the metallographic results the specific downward trend is clearly evident. The slope of the lines is the same within the experimental variations. As shown, the predicted and actual last layer to solidify differ by approximately 3 layers. Since the coil thickness is composed of multiple droplets, material flow may have an effect on heat diffusion and account for some of the discrepancy.

The integration zone is approximately 6 layers thick for our experiments.Since there is only a small zone of optimally integrated material, it is clear that temperature control is needed to build a fully integrated component. Due to the high deposition rates, arbitrary geometry, and thermal response time, substrate temperature control would be very difficult if not impossible. Thus a larger optimal integration window should be realizable by controlling the drop arrival temperature.Summary This work, though still continuing, has demonstrated several important findings. First, a novel calibration scheme for PDM is introduced. Second, we have shown that the bonding characteristics of the fabricated components can be predicted from a 1-D model with good accuracy.

First, a novel calibration scheme for PDM is introduced. Second, we have shown that the bonding characteristics of the fabricated components can be predicted from a 1-D model with good accuracy. Ongoing work is focused on extending the model predictions and experiment to aluminum components.Acknowledgments We gratefully acknowledge the generous funding from the National Science Foundation, grant number DMI-0070053, "High-Speed Droplet-Based Manufacturing of High-Precision Structural Components and Photonic BandGap Materials."The authors are also grateful for the assistance of Mr. Robert Smith and Mr. Tyler Pugh in the success of this project.References1 Jacobs, P. F., Rapid Prototyping and Manufacturing: Fundamentals of StereoLithography Published by the Society of Manufacturing Engineers, 19922 Orme, M., and Smith, R. “Enhanced Aluminum Properties by Means of Precise Droplet Deposition” ASME Journal of Manufacturing Science and Engineering, Vol.

AT TREATMENT EFFECTS ON MECHANICAL PROPERTIES OF WIRE ARC ADDITIVE MANUFACTURED Ti-6Al-4VNatalia Saiz*, Jonathan Pegues*, Shaun Whetten*, Andrew Kustas*, Tyler Chilson†Material, Physical, and Chemical Sciences Center, Sandia National Laboratories, Albuquerque, NM 87185, USA†ND Modernization & Future Systems, Sandia National Laboratories, Livermore, CA, USAAbstractDirected energy deposition (DED) is an attractive additive manufacturing (AM) processfor large structural components. The rapid solidification and layer-by-layer process associated with DED results in non-ideal microstructures, such as large grains with strong crystallographic textures, resulting in severe anisotropy and low ductility. Despite these challenges, DED has been identified as a potential solution for the manufacturing of near net shape Ti-6Al-4V preforms.

Despite these challenges, DED has been identified as a potential solution for the manufacturing of near net shape Ti-6Al-4V preforms. In this work, we explore several heat treatment processes, and their effects on tensile properties of wire arc additively manufactured (WAAM) Ti-6Al-4V. A high throughput tensile testing procedure was utilized to generate statistically relevant data sets related to each specific heat treatment and sample orientation. Results are discussed in the context of microstructural evolution and the resulting fracture behavior for each condition as compared to conventionally processed Ti-6Al-4V.IntroductionLarge format additive manufacturing (AM) using wire-based feedstock has becomeincreasingly popular for the manufacture of titanium alloy preforms reducing both waste and lead time associated with conventional casting and forging processes.

A common wire feed AM process is Wire Arc Additive Manufacturing (WAAM), which leverages typical welding energy sources such as MIG, TIG, plasma, or laser [1-4]. Wire directed energy deposition (W-DED) Ti- 6Al-4V often form large columnar prior beta (β) grains that encourage continuous alpha (α) grain formation along the beta grain boundaries. The size of the continuous alpha grains is mostly influenced by the cooling rate from the beta phase field with the highest cooling rates resulting in the smallest continuous alpha in highly anisotropic mechanical properties.The primary influence on strength for Ti-alloys is concentration of interstitial oxygenwhich can be of notable concern for welding AM processes that rely on shielding gasses to prevent significant oxidation during fabrication. The increase in oxygen and therefore strength is accompanied by severe reductions in ductility.

The increase in oxygen and therefore strength is accompanied by severe reductions in ductility. Alpha lamella width, grain size, and colony size also play an important role in Ti-6Al-4V mechanical properties. The slip length, which is governed by the size of the grain/colony or lamella width, results in increasing strength and ductility with decreasing size. As such, high cooling rates from the beta field are favored over slower cooling rates. Conversely, extreme cooling rates result in martensitic transformation1423resulting in acicular grains with high dislocation densities. These characteristics increase the strength of the alloy but have a deleterious impact on ductility and are therefore unfavorable for most applications.For W-DED Ti-6Al-4V the large columnar prior beta grains and associated continuousalpha are the primary concern.

These AM processes typically have more restrictions on deposition paths and lower cooling rates than powder-based AM processes, which is advantageous to continuous alpha formation and growth. Large continuous alpha grains result in early onset plastic deformation which mainly impact ductility. For columnar prior beta microstructures common in W-DED processing of Ti-6Al-4V, the directionality of the continuous alpha can result in severe anisotropy in ductility depending on how stress is applied. When the continuous alpha grains are subjected to a tensile opening stress mode, the deleterious effect on ductility is maximized [5]. This effect is observed when stresses are applied perpendicular to the build direction.

This effect is observed when stresses are applied perpendicular to the build direction. Post-processing thermos-mechanical heat treatments are therefore typically required to minimize these effects on material properties.This work explores two conventional post-process heat treatments for WAAM Ti-6Al-4Vto investigate their impact on critical microstructure characteristics and the resulting anisotropy of the mechanical properties.Experimental MethodsA billet of mill annealed Ti-6Al-4V was obtained from Bulk AM material which wasprocessed and rolled in the x-direction. Ti-6Al-4V wire feed stock was purchased from United States Welding Corporation with the compositions of 6.75% Aluminum, 4.5% Vanadium, and 0.15% Iron. WAAM specimens were created using 1/8” tungsten electrode on a wirefeed AM printer, print parameters are shown in Table 01.

WAAM specimens were created using 1/8” tungsten electrode on a wirefeed AM printer, print parameters are shown in Table 01. Specimens for both WAAM and mill annealed were extracted using wire EDM, dimensions are shown in Figure 01.Table 01: WAAM processing details.Wire Feeder Parameters: Delay Time: Feedrate:Hatch: Layer Height: Wire Feed Speed: 6Al-4V ELI Ti Wire:3.75 s 200 mm/min Table Speed infill – 150 mm/min perimeter 5 mm 1.25 mm 1350 mm/min Wire and Rod 0.045” diameter Al = 6.75% Fe = 0.15 % V = 4.5% Ti Bal.1424Specimens were extracted in three orientations related to the rolling and build direction asshown in Figure 01. Z-type specimens had their longitudinal axis perpendicular to the rolling direction in the xz-plane for the wrought material and parallel with the build direction in the xz- plane for the WAAM condition.

For the wrought material, the X-type specimens had their longitudinal axis parallel to the rolling direction within the xy-plane with Y-type specimens were perpendicular to the rolling direction also within the xy-plane. For the WAAM condition both X- type and Y-type specimens were perpendicular to the build direction with there axis aligned with their respective directions within the xy-plane.c)Figure 01: Direction of excised specimens from a) wrought Ti-6Al-4V, b) WAAM Ti-6Al-4V, and c) dimensions of wrought and WAAM Ti-6Al-4V high throughput tension specimens.Specimens were nondestructively inspected using an immersion ultrasonic method usinga Mistras Immersion tank with a 30 MHz 51 mm focal depth ultrasonic transducer probe. Immersion inspections were performed from the etched surface.

Immersion inspections were performed from the etched surface. The raster pattern used was done with a 50 µm resolution and at a height of 51 mm above the respective side being scanned.This work explores two standard heat treatments specific for Ti-6Al-4V which are summarized in Table 02. Heat treatment 01 (HT-01) is a standard beta anneal with over-aging to better match the mill annealed condition. The second heat treatment (HT-02) is an alpha anneal also followed by an over-aging procedure. Before heat treatment the samples were lightly etched to remove the brass recast layer deposited to the specimens from the brass EDM wire. Following cleaning the samples were subjected to the respective heat treatments in a vacuum furnace and1425quenched using Ar purge into the heating chamber. Both heat treatment schedules are shown in Figure 02.Table 02: Heat Treatment schedules for both LB-DED processes.HT IDHT-01 HT-021-Max Temp.

Both heat treatment schedules are shown in Figure 02.Table 02: Heat Treatment schedules for both LB-DED processes.HT IDHT-01 HT-021-Max Temp. (°C) 1050 9261-Soak (min) 60 601-QuenchArQ ArQ2-Max Temp. (°C) 725 7252-Soak (min) 120 1202-QuenchArC ArCFigure 02: Heat treatment schedules for a) beta anneal (HT-01, orange) b) alpha anneal (HT-02, blue), and c) overage anneal (black).Microstructural characterization was performed on a Zeiss Supra 55-VP FE-SEM equipped with Oxford Instruments X-Max SDD EDS and Symmetry EBSD detectors. EDS maps were collected at 20 kV across the entirety of the specimen build height, along the graded composition axis, to capture the full range of compositions, acquiring at least 106 counts per area. EBSD was performed at 20 kV for select regions to assess the grain and phase characteristics for all conditions.Oxygen analysis was performed on a LECO ONH 836 Interstitial analyzer.

EBSD was performed at 20 kV for select regions to assess the grain and phase characteristics for all conditions.Oxygen analysis was performed on a LECO ONH 836 Interstitial analyzer. All samples were placed in nickel capsules to aid in melting, then placed in a graphite crucible with graphite powder to convert any oxygen to carbon monoxide or carbon dioxide for analysis. Different quality control (QC) samples were measured to verify the analyzers were calibrated correctly. The QC’s were run after the calibration, during the run, and after the run. All the QC’s were well within the allowed range of the given value.All samples were tested using the 1B Mini Bionix frame with a 5000lbf load cell. Onecamera (FLIR 90 fps, 4.1 MPix) was used to capture images of samples during testing.

Onecamera (FLIR 90 fps, 4.1 MPix) was used to capture images of samples during testing. This camera captured images from the front of each tensile sample and had a 0.5x lens attachment with 12x zoom and a 1x adaptor.1426Experimental ResultsMill Annealed MicrostructuresThe mill annealed condition grain morphology and phase distribution is shown in Figure 03 showing the yz-plane perpendicular to the rolling direction. Figure 03(a) shows inverse pole figure (IPF) map illustrating the alignment of the grains in the -12-10 direction and partially recrystallized grains as indicated by the color gradient retained in many of the grains. Figure 03(b) shows the respective alpha + beta phases of the conventionally processed material. Approximately 5.9% of the beta phase is retained, as shown in blue.

Figure 03(b) shows the respective alpha + beta phases of the conventionally processed material. Approximately 5.9% of the beta phase is retained, as shown in blue. The average alpha grain size for the mill annealed condition is 5.32 µm.Figure 03: EBSD results of an excised Wrought Ti-6Al-4V sample showing a) inverse pole figure and b) phase maps.WAAM MicrostructureThe WAAM process results in high solidification rates and martensitic β-to-α‘transformation. This resulting microstructure can be observed in IPF map shown in Figure 04(a) as fine acicular grains. Compared to the mill annealed condition, Figure 04(b) reveals negligible retained beta for the as-built WAAM condition, <0.1%.

Compared to the mill annealed condition, Figure 04(b) reveals negligible retained beta for the as-built WAAM condition, <0.1%. The high cooling rates, which are directly correlated to the width of the acicular grains, result in an average grain width of 1.80 µm.1427Figure 04: EBSD results of an excised WAAM sample showing a) inverse pole figureand b) phase maps.The heat treatment effects on the resulting microstructure for the WAAM condition isshown in Figure 05. The alpha anneal stays below the beta transformation temperature (sub- transus) resulting in recrystallization of the martensitic α′ phase into a more typical alpha + beta microstructure for Ti-6Al-4V. The grains are a mixture of the lamellar and equiaxed grains as shown in Figure 05(a) with retained beta shown in Figure 05(b) as the blue regions mostly occurring at the grain boundaries of the alpha phase.

Contrastingly, the beta anneal heat treatment surpasses the beta transformation temperature (super-transus) then is rapidly quenched resulting in the martensitic β-to-α‘ transformation shown in Figure 05(c), with similar morphology to the as-built condition. This again results in the fine acicular microstructure and negligible retained beta as shown in Figure 05(d).1428Figure 05: SEM results of a heat treated WAAM Ti-6Al-4V sample showing a) alpha annealed condition with associated b) phase distribution and c) beta-annealed condition with associated d) phase distribution.Chemical CompositionsEDS was conducted on each material to document any major differences in chemicalcompositions since the wrought material and feedstocks originated from different sources. The results are tabulated in Table 03. The mill annealed condition Al composition was closer to 7 Wt.%, while the WAAM material was closer to 6 Wt.%.

The results are tabulated in Table 03. The mill annealed condition Al composition was closer to 7 Wt.%, while the WAAM material was closer to 6 Wt.%. Oxygen analysis showed the wrought material had a higher oxygen concentration compared to the WAAM produced material, shown in Figure 06. Furthermore, oxygen results demonstrated that the heat treatments resulted in minimal increases in oxygen.Table 03: Wrought and As-Built WAAM chemical compositions determined by EDS.WroughtPhase Al V FeAt. % 6.84 3.77 0.14As-Built WAAM At. % Phase 6.05 Al 3.68 V 0.12 Fe1429Figure 06: Oxygen Analysis for Wrought and WAAM As-Built, HT-01 and HT-02.PorosityFor the mill annealed condition there were no significant pores found in any of thespecimens. The ultrasonic inspection did identify significant pores in samples extracted in all orientations of the WAAM material.

The ultrasonic inspection did identify significant pores in samples extracted in all orientations of the WAAM material. In some cases, the pores were in the shoulder or grip section while other pores were in the gage section of the specimens. Seven WAAM samples displayed significant pores.Mechanical PropertiesThe stress-strain curves for the wrought and as-built WAAM conditions are shown inFigure 07 along with box and whisker charts comparing the tensile properties for each condition. The total range of the respective property for the wrought material is outlined by the dotted lines. Briefly, the strengths of the WAAM materials are noticeably lower than the wrought material while the ductility was lower on average but much closer to the range of the wrought material. In both cases some anisotropy was observed in yield stress and UTS. For the wrought condition the X-type specimen showed the highest average yield stress followed by Z-type and Y-type respectively.

In both cases some anisotropy was observed in yield stress and UTS. For the wrought condition the X-type specimen showed the highest average yield stress followed by Z-type and Y-type respectively. The box and whisker chart for yield stress suggests that the differences in each direction are significant as there is minimal overlap for each condition. The anisotropy in UTS for the wrought material was lower than yield stress with considerable overlap between the Y- and Z-type specimens. The X-type specimen, however, had the highest average strength which was determined to be significantly different from both Y- and Z-Type specimens. Interestingly the anisotropy in ductility did not appear to be significant though the averages for each type are noticeably different. There is considerable overlap between the box and whisker charts between all directions, suggesting these differences in averages are not significant.

There is considerable overlap between the box and whisker charts between all directions, suggesting these differences in averages are not significant. Conversely, the WAAM material showed lower strengths in Z-type specimens, both yield and UTS, but these were determined not to be significant as there is considerable overlap of the box and whisker data.1430Figure 07: Tensile properties for Wrought and As-Built WAAM. Yield Stress, Ultimate Tensile Stress, and Elongation to Failure for Wrought and WAAM As-Built.The stress-strain curves for the WAAM condition including all heat treatments are shown in the top images of Figure 08 with the respective box and whisker charts again provided below. All specimen directions are included in each graph. For the WAAM material both heat treatments resulted in significant decreases in ductility with the beta anneal (HT-01) resulting in a severe loss of ductility.

For the WAAM material both heat treatments resulted in significant decreases in ductility with the beta anneal (HT-01) resulting in a severe loss of ductility. The α anneal (HT-02) condition shows similar yield and UTS compared to the as-built condition while the beta annealed condition is so brittle that observations on the strength are difficult to determine.The highest yield average stress and UTS occur in the Y-type specimens for both heat treatments and as-built conditions, however, the difference compared to the other directions is negligible. For yield stress the difference from the highest average stress to the lowest average stress is 8%. Similarly, the UTS only had a difference of 6% between the highest and lowest values. The elongation to failure showed the highest anisotropy between averages at 19%.

Similarly, the UTS only had a difference of 6% between the highest and lowest values. The elongation to failure showed the highest anisotropy between averages at 19%. The box and whisker charts suggest these differences in strengths were not significant while the ductility for X-type and Z-type specimens from HT-02 do appear to be significantly different. Again, the beta annealed (HT-01) condition behaved in such a brittle manner that comparison of the mechanical properties for each direction are less reliable.1431Figure 08: Tensile properties for all a) WAAM As-Built, b) WAAM alpha, and c) WAAM beta anneal. Yield Stress, Ultimate Tensile Stress, and Elongation to Failure for WAAM As-Built, WAAM alpha and WAAM beta anneal.DiscussionOne distinct observation comparing the mill annealed conditions to WAAM conditionsare the higher strengths observed for the wrought material. This observation aligns with the observations made on the chemical analysis using EDS.

This observation aligns with the observations made on the chemical analysis using EDS. The WAAM process resulted in approximately 12% lower concentration of aluminum which contributes to the solid solution hardening behavior of Ti-6Al-4V. While printing the WAAM samples the composition of aluminum decreased due to the high energy input and aluminum having a lower boiling temperature compared to the other elements in Ti-6Al-4V. In addition, the wrought Ti-6Al-4V material was found to have a much higher oxygen content as compared to the WAAM produced material. As an interstitial, oxygen is a known strengthening mechanism for Ti-alloys and therefore the higher strength for the wrought material is mostly attributed to the higher oxygen content.The anisotropy for Ti-6Al-4V is often a major concern for additively manufacturedmaterials which is typically sensitive to grain boundary alpha that preferentially forms at the prior-beta grain boundaries.

Surprisingly, the lowest anisotropy was observed for the as-built WAAM material which suggests the cooling rate is sufficiently high to prevent large continuous alpha growth along these boundaries. The impact of the prior-beta microstructure and continuous alpha is still evident in the lower average ductility for the x and y-directions. Applying stress in these directions induces a tensile opening mode on the continuous alpha along the prior beta grains, resulting in early onset deformation and lower ductility. These differences in the as-built condition, however, are not considered to be significant.Interestingly, the alpha anneal heat treatment did show significant anisotropy between theductility in the X-type and Z-type directions, though there is considerable overlap with both directions compared to the Y-type data. This observation suggests that the sub-transus annealing1432step did promote continuous alpha growth along the prior-beta grains.

This observation suggests that the sub-transus annealing1432step did promote continuous alpha growth along the prior-beta grains. Figure 09 shows a representation of a continuous alpha grain for the alpha anneal condition. Continuous alpha results in a region where the slip length is much larger compared to the surrounding matrix. This size difference results in early onset plastic deformation when stresses are applied perpendicular to these continuous grains, resulting in the anisotropy observed.Figure 09: EBSD map of a continuous alpha grain along a prior beta grain resulting in increased anisotropy for the alpha anneal condition.Surprisingly, the WAAM material demonstrates the best material properties in the as-built condition. Anisotropy increased for the alpha anneal (HT-02) because of continuous alpha growth along prior-beta grains. This observation was also accompanied by lower average elongation to failure in all directions.

This observation was also accompanied by lower average elongation to failure in all directions. These observations demonstrate the important role the initial prior-beta microstructure plays in the mechanical properties of WAAM Ti-6Al-4V and how post-process heat treatments can magnify these effects. This also suggests that conventional heat treatment processes may not be appropriate for W-DED Ti-6Al-4V where columnar prior beta microstructures are present.ConclusionWAAM Ti-6Al-4V subjected to super-transus (HT-01) and sub-transus (HT-02) annealand overaged heat treatments were compared to an as-built and conventionally processed material.

Microstructural and mechanical properties in the three primary directions were analyzed and the following conclusions were made:WAAM material in the as-built condition showed more consistent mechanical properties across all directions compared to both post-process heat treatments.The as-built prior beta microstructure plays a dominant role in anisotropy with conventional post-processing heat treatments increasing these effects.1433Subsequent heat treatments for the WAAM material increased the oxygen content, however, the final content was still lower than the mill annealed material, resulting in lower strength.Nondestructive inspection for the WAAM condition suggests porosity does not significantly impact the mechanical performance.Results from this study demonstrate that the performance of the WAAM Ti-6Al-4V in the as-built condition is mostly consistent.

One note is that the high throughput testing utilized in this study provides a rapid means of comparison between various material conditions, however, the analysis is restricted to comparison of the miniature test specimens. To establish true design properties additional efforts are required that strictly follow established testing, and qualification standards.1434References[1] B. Baufeld, E. Brandl and O. Van der Biest, "Wire based additive layer manufacturing: Comparison of microstructure and mechanical properties of Ti–6Al–4V components fabricated by laser-beam deposition and shaped metal deposition," Journal of Materials Processing Technology, pp. 1146-1158, 2011.[2] F. Martina, J. Mehnen, S. Williams and P. Colegrove, "Investigation of the benefits ofplasma deposition for the additive layer manufacture of Ti–6Al–4V," Journal of Materials Processing Technology, pp. 1377-1386, 2012.[3] F. Wang, S. Williams and M.

1377-1386, 2012.[3] F. Wang, S. Williams and M. Rush, "Morphology investigation on direct current pulsed gastungsten arc welded additive layer manufactured Ti6Al4V alloy.," The International Journal of Advanced Manufacturing Technology,, pp. 597-603, 2011.[4] D. Yuan, S. Shao, C. Guo, F. Jiang and J. Wang, "Grain refining of Ti-6Al-4V alloyfabricated by laser and wire additive manufacturing assisted with ultrasonic vibration," Ultrason Sonochem,, 2021.[5] B. E. Carroll, T. A. Palmer and A. Beese, "Anisotropic tensile behavior of Ti–6Al–4V components fabricated with directed energy deposition additive manufacturing.," Acta Materialia, pp. 309-320, 2015.1435.

e Heterogeneous Compensation for the Infiltrative Error ofthe Binder Jetting Additive Manufacturing ProcessesXulong MA, Feng LIN, Lei ZHANGKey Laboratory for Advanced Materials Processing Technology, Ministry of Education of China Biomanufacturing and Rapid Forming Technology Key Laboratory of Beijing Rapid Forming Center, Department of Mechanical Engineering, Tsinghua University Beijing, 100084, P. R. ChinaAbstract: In binder jetting additive manufacturing, such as Three Dimensional Printing (3DP) and Patternless Casting Manufacturing (PCM, a process similar as ExOne and VoxelJet), the building error is mainly caused by the infiltration by the binder in the powder bed, and appear heterogeneous magnitude along different orientations because of the different infiltrating depth of the printed binder between the building direction and the binder printing direction.

Current methods to compensating these error are mostly based on the contour equidistant offset and the model shrinkage, which couldn’t deal with the heterogeneous infiltrative error. In this paper, we will propose a novel compensation method, in which the STL model will be counteracted heterogeneously in different directions to compensate the heterogeneous infiltrative distances of the binder in the powder. By this method, a sphere STL model will be transferred into an ellipsoid with variant axis length along different X/Y/Z directions. The method could greatly improve the dimensional accuracy of a series of additive manufacturing techniques which are based on the binder jetting onto powder bed.1. IntroductionAdditive manufacturing is a novel freeform fabrication mode, different with traditional removal fabrication, it is an accumulating manufacturing method from the top down driven by the three dimensional model.

Rapid Prototyping(RP), Solid Freeform Fabrication(SFF), Three Dimensional Printing(3DP), Additive Manufacturing(AM) are different names of this technique above. Nowadays, there are various kinds of AM technique, such as 3DP, PCM, DSPC, SLA, LOM, FDM, SLS, et al[1]. In the binder jetting additive manufacturing technology, including 3DP[2, 3], DSPC, PCM[4] et al, the binder is printed onto the powder bed to selectively bond the powder to form the prototype.In binder jetting AM technology, the droplet of binder is generated in a printhead, in which two different technologies are employed[5]: drop on demand(DOD) and continuous jetting. Continuous jetting continuously discharges a stream of fluid, which is broken into droplet by vibration. The technique is accomplished by selectively catching certain droplets as they pass through a collection equipment. On the other hand, in DOD technique, electrical pulses drive the piezo actuator in printhead to create separated droplet directly.

On the other hand, in DOD technique, electrical pulses drive the piezo actuator in printhead to create separated droplet directly. By integrating multiple piezo actuators as a nozzle array into a printhead, the building efficiency is increased greatly. In recent market, ExOne (Huntingdon, PA), Voxeljet (Augsburg, Germany) and Z-Corporation (Burlington, MA)[6] provide the various 3D printing systems based on DOD binder jetting technology and nozzle array printhead,200which are capable to manufacture the molds or sand models for metal casting.The feature size and fabrication rate of a part printed by the binder jetting AM technique are affected by the binder droplet size. The droplet size is the most important parameter for the balance between the building accuracy and fabrication rate. With the smaller droplets, the dimension accuracy becomes higher, but printing layer thickness becomes thinner and the fabrication time of a part will become longer.

With the smaller droplets, the dimension accuracy becomes higher, but printing layer thickness becomes thinner and the fabrication time of a part will become longer. If increasing the droplets and its infiltrative depth, the layer thickness could become thicker and the fabrication rate is showing to advantage, but the transverse infiltrative distance will increase either and the error caused by binder droplet infiltration turns out to be considerable. So if we can compensate the infiltrative error, the bigger droplet and less expensive printhead could be adopted. Furthermore, the binder jetting AM technique may become more reliable and efficient.With a DOD printhead integrated a nozzle array, a raster scan mode is usually applied, in which the printhead moves always fast in transverse direction, like X direction, and intermittently in longitudinal direction, like Y direction.

Since the anisotropic motion of printhead and gravity affect in vertical direction, like Z direction, it is noted that the droplet infiltrative distances become anisotropic and cause the error anisotropically as well, which is observed in the Patternless Casting Manufacturing (PCM) technique developed by Tsinghua University in 2002[7]. In this paper, the anisotropic infiltration error is investigated and a heterogeneous compensation method is presented.2. Dimension ErrorIn traditional binder jetting AM technology, a model with STL file format, which is a de facto standard widely used in CAD and AM field[8], is sliced along the direction of fabrication(Z axis) into a serial of discrete 2D contours. An error compensation is implemented in the directions of X/Y axes on two dimensional plane, the 2D layer contours take an equidistant shrink, and all are applied with an offset operation of the same distance.

In this method, that the Z axis error is not handled causes a significant error in Z direction. To compensate this error in Z direction, a model shrinks inward with the same distance before slicing operation. Once the sliced 2D contours are compensated with offset operation, the contours are converted to the bitmap images which can be recognized by the AM machine.A real infiltrative process of the printed binder appear heterogeneous magnitude along different orientations because of the different infiltrating distance in X/Y directions, the traditional compensation method does not deal with the difference of the heterogeneous error. This limitation will significantly affect the accuracy and quality of the fabrication, hindering the industrial applications of binder jetting AM technology.2.1 Infiltration ErrorA complete component fabricated by binder jetting AM technique can be viewed as a serial of single binder drop primitives stitched together to make a cohesive structure.

It is apparent that the size and shape of the primitive will ultimately determine the minimum feature size in the X, Y, and Z directions, as well as the surface finish of the component, all of which are critically important in making a high quality part.201(a) The infiltrative error (b) The primitive of a binder drop Figure 1. The phenomenon of infiltration of a binder drop in powdersThe actual infiltration of a binder drop in the powders is not equidistant along the X, Y, and Z directions. It is obviously different between the depth of Z direction and the distance of X or Y direction, while in fact it is also different between the depth of X direction and the depth of Y direction. So the binder infiltration has heterogeneous magnitude along different orientations. The binder drop primitive in the powders is definitely a half of ellipsoid as showed in Figure 1.

So the binder infiltration has heterogeneous magnitude along different orientations. The binder drop primitive in the powders is definitely a half of ellipsoid as showed in Figure 1. The ellipsoid’s equatorial radius and polar radius, a, b and c, are the infiltrative distance in X, Y, and Z directions, determined by the real situation. And it is relative with the physical property of binder and powders. The primitive is a form of ellipsoid with different equatorial radiuses and polar radius, usually a is approximate with b, but c is considerably different with a and b.2.2 ExperimentsIn this paper we used PCM technique to research the infiltrative error, so the binder is furan resin and the powder is the sand for foundry about 200 meshes. In PCM technique, the resin is injected onto the sand bed by the printhead, and the catalyst is mixed into the sand in advance. So after removing the un-printed sand, the part attained will be solidified through chemical reaction.

So after removing the un-printed sand, the part attained will be solidified through chemical reaction. To investigate the different distance of binder infiltrative error in X, Y, and Z direction, an experiment part with a few of geometric feature is designed as Figure 3. In Figure 3, the features named A1 to A5 are simple geometric body, just cubes and cylinders; the features named B1 to B10 are gaps with the width from 0.4 to 6.4 millimeters. We fabricated two parts, part A and part B, the part A has the original dimension, while the part B has an offset of every layer contours with 1.2 millimeters. The offset value is chosen tentatively.(a) Length of X/Y/Z(b) Gap Width of X/YFigure 3. The features on the experiment part.202Figure 4. The experiment part A and part B manufactured by PCM technique.We measured the two parts’ feature size, features of A1 to A5, in X, Y, and Z directions, in Figure 4.

The experiment part A and part B manufactured by PCM technique.We measured the two parts’ feature size, features of A1 to A5, in X, Y, and Z directions, in Figure 4. Compared with the designed dimension size, the actual size distribution and error distribution are showed in a histogram in Figure 6. The most feature size of part A has a positive error value in X and Y direction, while the part B has a negative error value. It explains two points that there is an infiltrative distance in both X and Y directions and the part B with the contour offset of 1.2 millimeters has an excessive compensation. In Z direction, part A and B both have a positive error value, because the traditional compensation method does not handle the errors in Z directions. In table 1, the average feature lengths in X, Y, and Z direction are calculated. The average error is also figured out, listing in table 2.

In table 1, the average feature lengths in X, Y, and Z direction are calculated. The average error is also figured out, listing in table 2. Finally we get a consequence of different error sizes in X, Y, and Z directions, a proof of the heterogeneous infiltrative error in PCM technique on behalf of binder jetting AM technology. In our case, there is a difference of 0.4 millimeters between X error and Y error in a good agreement of both part A and B. The error value in Z direction presents a proximity for a lack of compensation in both part A and B.Table 1. The dimensions of the experiment parts (mm)XA1-A4 YZA5X-Outer Y-Outer X-Inner Y-Inner ZTarget Part A 15.40 15.18 20.40 Part B 14.33 14.12 20.6915152040 40.27 39.2040 39.89 38.8430 30.45 29.3630 29.93 28.85Table 2.

The errors on direction of X/Y/Z (mm) Delta Z 0.42 0.51Delta X 0.38 -0.69Delta Y 0.09 -0.98Part A Part B20330 30.51 30.80Feature Dimension Error Distribution1.50 1.00 0.50 0.00 -0.50 -1.00 -1.50XYXYXYZXYZXYZXYZ OD ID OD ID ZXYZA1A2A3A4A5AveragePart APart BFigure 6. The dimension errors distribution of the features named A1 to A5The gaps with a serial width from 0.4 to 6.4 millimeters, parallel with Z direction, are designed to inspect the manufacturing capability of the PCM machine. Since having an excessive offset compensation, the part B has a larger gap width than the target size. While the part A has an insufficient gap width distinctly for a lack of compensation. A phenomenon of gap disappearance is observed in part A, shown in Figure 9, illustrating that there is an infiltrative scope not less than 0.4 mm. The gap widths of X direction and Y direction are listed in table 3.

The gap widths of X direction and Y direction are listed in table 3. There is also a difference about 0.3 millimeters between X and Y directions in gap widths statistics.Table 3. The dimensions of the test part marked with B (mm)XYXYYDXDYTarget Part A Part B6.44.257.503.21.471.471.60.790.790.80.4NONE NONE1.280.6601.480.6301.730.31Gap Width Error8.00 7.00 6.00 5.00 4.00 3.00 2.00 1.00 0.00 -1.00 -2.00 -3.00Part BPart AB1B2B3B4B5B6B7B8B9B10Figure 8. The gap width error distribution of the features named B1 to B10204(a) B7 width=0.8 mm (b) B8 width=8 mm(c) B9 width=0.4 mm (d) B10 width=0.4mm Figure 9. The disappeared gap in experiment part B3. Compensation MethodSince the heterogeneous infiltrative error is confirmed, we propose a new compensation method to make up the deficiency of the traditional 2D method.

Compensation MethodSince the heterogeneous infiltrative error is confirmed, we propose a new compensation method to make up the deficiency of the traditional 2D method. By the new heterogeneous compensation method, a sphere STL model will be transferred into an ellipsoid with variant axis length along different X/Y/Z directions. The method could greatly improve the dimensional accuracy of a series of binder jetting AM techniques. 3.1 Heterogeneous CompensationThe new compensation method adopts the original STL model as data source to be dealt with, every triangular facet will be offset a distance which is calculated by its normal vector and infiltration vector. The offset distance varies according to the facet’s position and direction, so the STL model will be counteracted heterogeneously in different directions to compensate the heterogeneous infiltrative distances of the binder in the powder.The new compensation method is applied before slicing in traditional processing.

After compensation, the new STL model will be used to the next processing without a compensation any more. The concept of line width in traditional compensation method is disappearing, and independent compensation in Z direction is also not needed again. Once the bitmap image is created, it can be directly used to control the AM machine.The new infiltrative compensation method includes a few steps as follows: (1) Construct the topological triangular mesh from the original STL data. Transformation of facets or vertices will not be done in this processing, it just find the topological information205among vertices, edges and facets which is used to guarantee the validity of the STL model for transformation in next step.(2) Read a triangular facet one by one into program, and calculate each one’s offset distance which depends on the normal vector and infiltrative vector.

The compensating offset value is defined as distance between the envelope plane formed by the stacked half infiltrative ellipsoid and the plane formed by the all ellipsoids’ center points, just shown as Figure 10.Figure 10. A sketch view of the compensating offset(3) Read the vertex into program one by one, calculate each one’s offset point through the offset distances of the triangular facets adjacent to the current vertex. The transformation based on vertices can keep the topological structure of all elements in STL model spontaneously. Of course, there may be no solution vertex satisfying the offset distances of adjacent facets, thus a vertex fission technique can be used to find a compromised solution with several vertices.

Of course, there may be no solution vertex satisfying the offset distances of adjacent facets, thus a vertex fission technique can be used to find a compromised solution with several vertices. There maybe are facets not a triangle created in this processing, this situation does not satisfy the rules of a STL model, so the triangulation is applied to the facets that does not match the rules.(4) Repeat the step 2 and step 3, until all facets and vertices are processed. (5) Checkout the validity of the STL data, including the correction of topological information and the correction of the facet normal vector. Then a new valid compensated STL model is gained.(6) Furthermore, the next steps of data processing, including slicing STL model, do not need a compensation any more.

Then a new valid compensated STL model is gained.(6) Furthermore, the next steps of data processing, including slicing STL model, do not need a compensation any more. Thus the concept of line width disappears, and the data can be used to create a bitmap image for the AM machine directly.3.2 Case StudyThe infiltrative compensation method described in this paper has been implemented simply using a program language of C++. The compiler is gcc-4.7.3, the computer OS is Ubuntu 12.04 64bit. Several models are tested to see the performance of the new compensation method shown in Figure 12.206(a)A jaw model of STL with compensation(b)A factory part of STL with compensation Figure 12. The examples of model compensation4. ConclusionThe new infiltrative compensation method changed the data processing mode for compensation.

terogeneous Material Modeling with Distance FieldsVadim Shapiro(cid:3) Spatial Automation Laboratory y University of Wisconsin-MadisonArpan BiswasIgor TsukanovAbstractWe propose a universal approach to the outstanding problem of computer modeling of con- tinuously varying distributions of material properties satisfying prescribed material quantities and rates on a (cid:2)nite collection of geometric features. The central notion is a parameterization of the shape’s interior by distances from the material features - either exactly or approximately; this parameterization supports speci(cid:2)cation, interpolation, and optimization of desired mate- rial distributions in a systematic and controlled fashion.

The task is non-trivial because the modern geometric and solid modeling technology has been developed under the assumptions of material homogeneity.Informally, we can summarize the material modeling problem as follows: given a geometric representation of solid and/or collection of material features with known material properties, con- struct one or more material functions, subject to some given constraints (design, manufacturing, etc.) Each material function represents some material property that varies, usually continuously, from a point to point throughout the space, including the boundary and the interior of the solid.100% SiC80% SiC 20% DiamondFraction of SiCFunctionally Graded MaterialDiamond ChipFraction of diamond(a) Diamond cutter tool consists of several(b) Fractions of each material vary throughoutdistinct material regionsthe solid but always add to 1Figure 1: A typical material modeling problem requires construction of material property functions that interpolate known material features subject to speci(cid:2)ed constraints and physical lawsConsider a simple but typical example that is depicted in Figure 1(a) and discussed in [6].Three material features are identi(cid:2)able in the above example: the diamond chip and the two solid subsets of the shank (block region with 100% SiC, prismatic region with 80% SiC and 20% diamond,) where material properties are known.

Two material functions need to be constructed (one for SiC and another one for diamond), subject to the following constraints: continuous interpolation of the material fractions on each of the features, fractions must add to 1, and the rate of material change is linear in the distance from each feature. In a more general case, material features can be of any dimension, shape or topology; the material distributions may be given as known continuous func- tions f (x; y; z); the interpolation laws may include arbitrary weights and be subject to algebraic, differential, or integral constraints.The recent surveys [19, 17] describe the technical challenges of this task and compare several known approaches from the perspective of applications, namely design, analysis, manufacturing process planning, and data exchange. To summarize:(cid:15) Material functions should match as exactly as possible their speci(cid:2)ed behaviors on materialfeatures.

Many of such constraints are application speci(cid:2)c; for example, when the solid is composed from several materials, individual material functions represent a fraction of volume occupied by a particular material and must add to 1 at any point in the solid.(cid:15) The representation of material properties must be compatible with current or proposed stan- dards for geometric modeling representations as described in ISO 10033 standard [5, 15]. This is essential to support exchange data between design, analysis and manufacturing pro- cess plan domain.With the exception of the heuristic procedural method in [16], all proposed approaches to material modeling require some form of spatial discretization of the solid’s interior. These in- clude voxel-based [3, 25], (cid:2)nite-element based [18, 20], mesh-based [12, 8], and set-based [10, 9], or layer-based [24] schemes.

These in- clude voxel-based [3, 25], (cid:2)nite-element based [18, 20], mesh-based [12, 8], and set-based [10, 9], or layer-based [24] schemes. Such discretizations amount to representation conversions that are expensive to compute and lead to many complications. First of all, all discretization methods introduce errors because they must approximate the geometry of solids and material features, as well as prescribed material properties. Secondly, the ability to satisfy the constraints and to as- sure smoothness properties places substantial restrictions on the types of allowed discretizations and approximations. Thirdly, modifying and controling material models becomes next to impossi- ble, because every change may require recomputing both discretization and approximation of the material functions. Last, but not least, such representations are awkward for data exchange and standardization due to errors, approximations, and large size.

Last, but not least, such representations are awkward for data exchange and standardization due to errors, approximations, and large size. Comparing these dif(cid:2)culties to the list of the requirements above, it seems clear that approaches to material modeling based on spatial discretizations are not likely to provide the required level of support. The heuristic approach in [16] offers no guarantees but corresponds to a special restricted case of material modeling in our framework [2].1.2 Distance is the keyMuch of the existing literature on material modeling points in the direction of one natural param- eter: the distance from a material feature. For design purposes, it is convenient and intuitive to specify how material composition changes as a function of the distance from the material fea- ture (recall the example in Figure 1).

For design purposes, it is convenient and intuitive to specify how material composition changes as a function of the distance from the material fea- ture (recall the example in Figure 1). Capabilities of manufacturing processes for functionally- graded materials are also commonly described by their ability to modify material as functions of distance (determined either analytically or experimentally) [13, 7]. The most common types of material functions constructed by methods based on spatial discretization appear to be either the Euclidean distance function, weighted distance functions or simple functions of a distance function [4, 25, 24, 12, 8].However, two possible dif(cid:2)culties may arise in relying on distance (cid:2)elds: computational cost and loss of differentiability at equidistant points.

Computing the distance from a point to geomet- ric primitives (typically a curve or surface) usually requires a numerical iterative procedure [14], which may result in a high computation cost. In the context of material modeling, the lack of smoothness in a material function constructed as a function of distance will result in stress con- centrations or other undesirable singularities. We propose to eliminate both of these limitations of307the exact distance (cid:2)elds by replacing them with various smooth approximations, while preserving most of the attractive properties of the distance (cid:2)elds.

In particular, we can replace the exact dis- tance (cid:2)elds with their approximations constructed by the theory of R(cid:0)functions [21] (See Figure 2).(a)(b)(c)Figure 2: (a) The exact distance (cid:2)eld of an S-shaped 4th-order B-spline curve; (b) An approximate distance (cid:2)eld for the same curve normalized to the (cid:2)rst order; (c) A section of an approximate distance (cid:2)eld for a 4th-order B-spline surface normalized to the second order.In the following section we describe with examples how material functions can be constructed and controlled with the distance (cid:2)elds of a single and multiple material features in our modeling scheme. We often used two-dimensional examples for the explanation because they are much easier to understand and visualize. However, all constructions and techniques apply to three- dimensional material modeling problems without any changes.

However, all constructions and techniques apply to three- dimensional material modeling problems without any changes. Composition of several materials and other vector-valued material properties are modeled as vector-valued material functions in Section 3. A more detailed discussion of our approach, including theoretical details and many additional examples, can be found in [2].2 Single and Multiple Material FeaturesThe simplest problem of material modeling involves a single material feature. The material func- tion P is constructed and controlled by specifying the value and derivatives at the feature. Using the distance function of the material feature, value and derivatives of P , speci(cid:2)ed at the material feature, is extended to the whole domain. Such material function construction method relies on the generalized Taylor series expansion by powers of a distance (cid:2)eld of the feature and we showed that any material function can be constructed with our approach [21, 2].

The interpolation308(a)(b)(c)Figure 3: (a) Material function e((cid:0)1:5u) is prescribed in terms of distance u to the S-shaped feature. (b) Non-constant material function on a feature with derivative control: material function ye((cid:0)1:5u) and gradient -4 is speci(cid:2)ed on the feature. (c)Inverse distance interpolation of two material func- tions with derivative control: P 1 = 1 (cid:0) 1:5u (gradient speci(cid:2)ed at the curve is -1.5) is associated with the S-curve and P 2 = 2 is associated with the circular ring.technique, which is employed to construct the single material function, is known as inverse dis- tance weighting or as Shephard’s method when used for scattered data interpolation [23, 11, 22]. The underlying notion of the interpolation method is that the in(cid:3)uence of a feature decreases in proportion to the distance from the feature.

The underlying notion of the interpolation method is that the in(cid:3)uence of a feature decreases in proportion to the distance from the feature. Such interpolation technique does not require spatial discretization of the domain and applies to material features of arbitrary shape, topology and di- mension. It also preserves exactly the values and derivatives speci(cid:2)ed on each material feature. The material distribution function in Figure 3(c) was constructed with two material features: the ’S’-shaped curve and an annulus region. The function speci(cid:2)ed at the ’S’ curve was (1 (cid:0) 1:5u), where u is the distance function of the ’S’ curve. The constant material function 2 was speci(cid:2)ed at the annulus feature.(a)(b)Figure 4: Material modeling with differential constraint. (a) Three-dimensional material modeling problem requires constructing a material function for the solid from the material functions asso- ciated with the two features as shown.

(a) Three-dimensional material modeling problem requires constructing a material function for the solid from the material functions asso- ciated with the two features as shown. (b) Constructed material distribution constrained to satisfy Laplaces equation.The formulation based on generalized Taylor series expansion of the representation scheme309also makes possible to control material function with differential, integral, analytic and explicit constraints [2]. For example, Figure 4(b) shows the speci(cid:2)ed material function in Figure 4(a) constrained to approximate the Laplaces equation r2P 2 = 0.

For example, Figure 4(b) shows the speci(cid:2)ed material function in Figure 4(a) constrained to approximate the Laplaces equation r2P 2 = 0. Material functions for two material features are speci(cid:2)ed: P 1 = 1 (cid:0) u2 1=2 for the material feature de(cid:2)ned as the union of all vertical faces (circular hole and four vertical faces of the cube), and P 2 = 0:1e(cid:0)u2 2 for the spiral canal surface through the interior of the solid.3 Vector Valued Material PropertiesThe proposed approach to material modeling extends directly to a more general case where a ma- terial property is a vector valued function. Common examples of such properties include material anisotropic grain orientation represented by a vector (cid:2)eld, material composition represented by a vector of volume fractions, and vector of varying shape inclusion parameters in microstructure models. These and other vector and tensor valued properties of materials are discussed and com- pared in [9].

These and other vector and tensor valued properties of materials are discussed and com- pared in [9]. Each scalar component function can be treated independently using the techniques of section 2, but the component functions are also constrained to satisfy additional conditions. For example, when composition of several materials is represented, each component function P 2;j(p); j = 1; 2; : : : ; m models a fraction of the total volume, and the sum of all components P 2;j = 1 at every point of space.

In [2], we showed that this constraint on volumet-m P j=1must beric factions can be maintained automatically using generalized inverse distance interpolation of the material feature functions P i;j satisfying the partition of unity condition, where P i;j is the jth material fraction function associated with the ith feature.For example, in the case of diamond cutter example of Figure 1, constant fractions of two materials, SiC and diamond, are prescribed on three features: P 1;1 = 1:0; P 1;2 = 0:0; P 2;1 = 0:8; P 2;2 = 0:2 P 3;1 = 0:0; P 2;2 = 1:0. These constant fraction functions are interpolated into two material fraction functions: P 2;1 for SiC and P 2;2 for the diamond. Each function was constructed using trans(cid:2)nite interpolation with inverse distance weights as shown in Figure 1(b).

Each function was constructed using trans(cid:2)nite interpolation with inverse distance weights as shown in Figure 1(b). By construction, the two material functions add to 1 at all points of the cutter, simply because each feature was prescribed fractions of materials that add up to 1.4 ConclusionsIn the context of material modeling, the distance (cid:2)elds provide a natural parameterization of the space and allow formulation and solution of the material modeling problem in a manner that covers most practical situations in modeling, design, and manufacturing. The theoretical completeness of the formulation guarantees that any and all material functions may be represented to a desired precision as functions of the distance (cid:2)elds associated with material features [2].

We conjecture that the practical methods described in this paper are suf(cid:2)cient for handling most applications involving heterogeneous materials.This paper considered relatively simple constraints on material properties. Additional research is needed on material functions to represent general tensorial properties, possibly anisotropic, with or without symmetry, and perhaps with periodicity such as found in composite materials. We al- ready observed that material modeling can be viewed as a particular kind of a boundary value problem. One could also argue that the material modeling is a more general problem, because physical (cid:2)elds (temperature, stress, electric charge, etc.) are essentially material functions that310are de(cid:2)ned by boundary conditions and constrained by suitable differential equations. A natural extension of our work is combining material modeling and engineering analysis within the same meshfree computational framework.

A natural extension of our work is combining material modeling and engineering analysis within the same meshfree computational framework. In fact, the same distance (cid:2)elds provide a natural represen- tation for any number of distinct physical (cid:2)elds over the same geometric domain, suggesting a possible approach to handling multi-physics problems. On the other hand, geometric domain itself may be also de(cid:2)ned implicitly by a threshold of the constructed material density function[1]. Sen- sitivity analysis with respect to either geometric parameters or material changes, as well as shape and material optimization, can be performed in a meshfree manner using the same computational utilities and avoiding the usual dif(cid:2)culties with (re)meshing [27].References[1] M. Bendsoe and N. Kikuchi. Generating optimal topologies in structural design using homog- enization method. Computer Methods in Applied Mechanics and Engineering, 71:197(cid:150)224, 1988.[2] A.

Bendsoe and N. Kikuchi. Generating optimal topologies in structural design using homog- enization method. Computer Methods in Applied Mechanics and Engineering, 71:197(cid:150)224, 1988.[2] A. Biswas, V. Shapiro, and I. Tsukanov. Heterogeneous material modeling with dis- tance (cid:2)elds. Technical Report SAL 2002-4, Spatial Automation Laboratory, University of Wisconsin-Madison, 1513 University Avenue, Madison, WI 53706, June 2002.[3] V. Chandru, S. Manohar, and C. Prakash. Voxel-based modeling for layered manufacturing.IEEE Computer Graphics and Applications, 27:42(cid:150)47, 1995.[4] W. Cho, E. M. Sachs, N. M. Patrikalakis, M. Cima, H. Liu, J. Serdy, and C. C. Stratton. In Proceedings of the 2002 NSF Local composition control in solid freeform fabrication. Design, Service and Manufacturing Grantees and Research Conference, San Juan, Puerto Rico, January 2002.[5] International Organization for Standardization.

Design, Service and Manufacturing Grantees and Research Conference, San Juan, Puerto Rico, January 2002.[5] International Organization for Standardization. Industrial automation systems and integrationproduct data representation and exchange. Technical Report part 10303, 1994.[6] Asish Ghosh, Yoshinari Miyamoto, and Ivan Reimanis. Functionally graded materials: Man-ufacture, properties and applications. Ceramic Transactions, 76, July 1997.[7] J. B. Holt, Koizumi M., Hirai T., and Munir Z. A., editors. Ceramic Transactions, volume 34.The American Ceramic Society, Westerville, Ohio, 1993.[8] T. R. Jackson. Analysis of Functionally Graded Material Object Representation Methods.PhD thesis, MIT, Ocean Engineering Department, June 2000.[9] V. Kumar, D. Burns, D. Dutta, and C. Hoffmann. A framework for object modeling.Computer-Aided Design, 31:541(cid:150)556, 1999.[10] V. Kumar and D. Dutta. An approach to modeling and representation of heterogeneous ob-jects.

K. Siu and S. T. Tan. Modeling the material grading and structures of heterogeneousobjects for layered manufacturing. Computer-Aided Design, 34:705(cid:150)716, September 2002.[25] Y. K. Siu and S. T. Tan. Source-based heterogeneous solid modeling. Computer-Aided De-sign, 34:41(cid:150)55, January 2002.[26] S. Suresh and A. Mortensen. Fundamentals of Functinally Graded Materials. The UniversityPress, UK, 1997.[27] I. Tsukanov and V. Shapiro. The architecture of SAGE - a meshfree system based on RFM.Engineering with Computers, accepted for publication, 2002.312In In.

terogeneous Solids: Possible Representation.SchemesStephane M. Morvan smorvan@ces.clemson.eduGeorges M. Fadel gfadel@ces.clemson.eduClemson Research in EngineyringDesign and Optimization Dept. of l\4.echanicalpngineering Clemson University, Clemson, SC 29634-0921AbstractSolid freeform fabrication processes allow parts to be built with accuracy and mechanical integrity, permittingthem to be used in tooling or fOrlnandfit applications. There is already a need form~lti..color Parts. for surgical applications, which will eventually lead to. multi-material RP .machines.WhetherJor on the spot color deposition or for functionally tailored multiple materials parts,.RPmachines with such capabilities are becoming available. They will eventually lead to the trtiepromise of Solid Freeform Fabrication: a system that can build a functional mechanism without assembly, and from multiple materials.

They will eventually lead to the trtiepromise of Solid Freeform Fabrication: a system that can build a functional mechanism without assembly, and from multiple materials. This paper is aimed at understanding the new challenges raised from representing solids whose material distribution is changing gradually from one material to another (HC), and those made of a collection .of discrete materials (HD). Several representation schemes are reviewed and critiqued. Techniques borrowed from medical imaging and geoscience modeling are used to better understand the modeling of heterogeneous and gradient solids, from a geometric standpoint.Problem Statementimplicitly making the assumption that the interiofof the solid is made of a single, homogeneous material. Recent progress in commercial RP solutions [1] will allow parts to be built with localized mechanical properties.

Recent progress in commercial RP solutions [1] will allow parts to be built with localized mechanical properties. Thus, as more efforts are directed· towards material optimization techniques [2],[3] there isa need to shift the focus of a solid representation scheme from the geometry to a .more abstract attribute..oriented representation scheme. In such a representation, the geometry of an object as it is conventionally known, would become the. residual of all the spaces spanned by the attributes under thermal properties, mechanical characteristics, electrical conductivity and density.Current solid modeling technologies model an object using its boundaries,consideration. Exampleattributesare:2 Rapid PrototypingThe enabling technology behind the creation of heterogeneous objects has to be able to output an artifact representing an object designed with a computer. RapidPrototyping (RP) seems to be the medium of choice in such applications. RPembodies digital. designs into.

RapidPrototyping (RP) seems to be the medium of choice in such applications. RPembodies digital. designs into. physical mockups in a matter of hours. The techniques behind the RP process rely on the ability to add (and/or in some casesremove~, material at anyarbitraryloqatipn.inspGlce.. Therearetwogeneral classes .. of RP processes (even though combinations do exist): .additive-based processes and removal..based processes. The distinction between the processes depends on whetherthe process adds.material to reach the final shape ofthe.artifact, orifitre1110vesmateriaLfromabulkto.reach187that same shape. The following two sections review those two important families of processes. This review is by all means non-exhaustive, and is here only to· briefly outline the principles behind them.

The following two sections review those two important families of processes. This review is by all means non-exhaustive, and is here only to· briefly outline the principles behind them. Further description of RP processes is proposed in [4].Additive Processesto be manufactured is broken down.in layers and the part is built from bottom to top by stacking cross sections of the object. Each'layerismanufactured byway of an additive process, one at a time andAdditive processes involve the following or a variation thereof:the objecta setorder, then the object isshiftedd9Wtl to allow the next layer to be added.Some of the enabling technologies. behind additive processes are based on the hardening of a photo-polymer by a laser (SLA) or by ultra-violet floodlights (SOC), or the cutting of an adhesive paper (LOM). Another variation, in powder-based processes, is to sinter a bin of finely grained powder (SLS) [5].

Another variation, in powder-based processes, is to sinter a bin of finely grained powder (SLS) [5]. Some processes also rely on a thermoplastic plotting system, akin to an inkjet plotter to deposit the material (FDM, .3Dprinters).• Many .of those processes require additional support structures to maintain the artifacts' structural integrity throughout the building stage, thus requiring the additional step of part cleanup.Some RP makers use a color-coded support material whose removal does not risk obliterating a feature from the part. In some other cases, the technological choice of the process prevents from using a secondary· material, and the support structures and the part both share the same material, with the exception being that the support structure is attached to the part only temporary by a breakable junction.2.2 Removal ProcessesThis class of processes is more on a par with traditional manufacturing technologies.

Essentially, it involves milling, grinding or cutting material from a bulk of raw material. Though this technique certainly has more limitations than the additive processes (recessed pockets are one of them), it does provide a valid alternative to obtain mechanical parts with properties and finish that are often unattainable with additive processes.The bulk material is typically made of a ductile material, which can be cut at a great rate of speed to quicken the process. Example materials are aluminum, steel, Styrofoam and balsa. It is obvious that for manufacturing reasons, the bulk material should be made of a single material, otherwise the rate of cutting will vary from one material to the next, resulting in uneven surfaces, or rips at the interface between the two materials.3Solid Modeling Representation SchemesTraditional mechanical design encompasses several tasks and disciplines.

For instance, designing a car's body involves considerations of structural soundness, ergonomics, aerodynamics, vehicle dynamics, aesthetics and after-life disposability. In addition, modem engineering often professes some form of optimality which insures that a designed product responds ideally to a set of constraints, both from the consumers' and from the manufacturer's perspectives.It is evident that today's design cycles are complex, demanding and expensive, and that the interplay of different disciplines requires rugged tools to assist the designers and to supportcollaborative and concurrent work. The widespread use of computers has greatly contributed to reduce the time required to perform the aforementioned activities, leading to a design cycle nearly 100% digital (i.e.

The widespread use of computers has greatly contributed to reduce the time required to perform the aforementioned activities, leading to a design cycle nearly 100% digital (i.e. performed entirely on computers).Ideally, the enabling technology behind a digital design cycle is made of tightly integrated software tools revolving around a central application called a CAD (Computer Aided Design) System. Most CAD Systems are mature software products, sometimes resulting from a decade of in a research or more. Behind all representation allowing the modifications, simulations and variations encountered during the design phase. Since mechanical design has always been concerned with the dimensioning of parts, a natural choice for this representation is based on the geometric features of a product and its dimensions.

Since mechanical design has always been concerned with the dimensioning of parts, a natural choice for this representation is based on the geometric features of a product and its dimensions. This is referred to as solid modeling, and the next section will briefly review the f(~presentation concepts behind the various modelers found in CAD Systems today.these CAD Systems is the ability to portray a product3.1 Boundary Representation Schemes (B-Rep.)The B-Rep is a surface boundary representation where solids are defined by a list of their enclosing surface boundaries. This representation originated from the definition of a polyhedra: it is made from polygons, which are made from triangles, which in turn are made from edges and lastly edges are made from vertices. The representation was then further expanded to include arbitrary representations (Beziers patches or NURBS surfaces) for the enclosing surfaces.

The representation was then further expanded to include arbitrary representations (Beziers patches or NURBS surfaces) for the enclosing surfaces. The surface boundaries are oriented such that given any surface point, the solid interior can be easily determined. For instance, a pyramid with a square base is made of four triangles and a square (Figure 1). This representation is notably good at representing most solids and especially the details embedded in complex shapes, though a known drawback of the B-Rep is the difficulty to verify whether or not a solid is closed or topologically valid.This difficulty emanates from the task of inferring from the enclosing surfaces (typically surfaces described by two parameters) that a valid, closed 3D entity is defined. A solid is closed when the enclosing boundaries form a closed/fixed volume. A solid is valid if it does not self intersect.

A solid is closed when the enclosing boundaries form a closed/fixed volume. A solid is valid if it does not self intersect. For a more rigorous description of a B-Rep solid topology and its problems, refer to [6].L7Figure 1: A Boundary Representation for a Pyramid with a Square Base.1893.2 Constructive Solid Geometry (CSG)The GSG· representation [7] defines solids as being made from three-dimensional volumetric primitives (blocks, cylinder, spheres, cones ... ) associated together using binary operators {intersection, union, subtraction) to form a binary tree (CSG tree) whose end result (the topmost node) is the desired solid. The binary operator requires two arguments to yield a result. A binary tr~.e [8] is defined recursively (up to a sufficient depth) as being constituted by a root node and two binary trees (which in tum are made of two nodes and four binary trees). In the case of a CSG tree,.

In the case of a CSG tree,. the end leaves are primitive solids and the topmost root node is the desired product, the intermediate nodes being the results of intennediate operations.Creating complex solids by combining simple primitives is intrinsically appealing when the need to l11anufacture a product appears, as some operators are homeomorphic to material· removal processes: for example, drilling a hole in a block is similar to subtracting a cylinder from that same block.Inherently, since the operators applied on the primitives are isomorphic, we are generally guaranteed that the resulting solids are both closed and valid. The cost of having a solid defined as being made from other interrelated solids is associated with the complexity required to manage the CSG tree which may need to be entirely recomputed when updates are made to the model.

Also, displaying the solid requires it to be converted into simple display primitives such as triangles, a non-trivial operation.It should be noted that most modern GSG solid modeling packages do rely at some point on a mixed representation of solids (mixed between B-Rep and GSG), as GSG lacks the capability to perform local operations such as fillet, a typical operation in B-Rep based solid modeling packages.

Overall, the GSG representation does an excellent job at representing objects, however, as object complexity grows, the traversal of the binary tree becomes increasingly computationally intensive often resulting in long updates for minor changes.3.3Set Theoryfollowing cOnsideration: "S is a solid which can be defined as follows, given a boundary B defined by an implicit equation f(x,y,z) = 0, M(XM,YM,ZM) is a point such that:This mathematicalrepresentation principle liesalmost entirely in thef(xM,YM,zM) < 0 => the point is outside Sf(xM,YM,zM) = 0 => the point is on the border of S (=B)f(xM,YM,zM) > 0 => the point is inside SThe set of all points inside S and on its border defines the solid. Unfortunately, the strict nature of the membership of a point M in the set (it is either in, on or out of the solid) precludes polyphase objects or gas mixtures to be represented accurately.

Unfortunately, the strict nature of the membership of a point M in the set (it is either in, on or out of the solid) precludes polyphase objects or gas mixtures to be represented accurately. The need to model these objects is now omnipresent: whether it is for the Finite Element analysis of an engine's explosion cycle or for a feature animation film, there is indeed a need to represent objects completely (which then raises the correct definition of an object).190Used initially for manufacturing, this set-based representation for solids has been outgrown by the demands put on it. An example application for this repr~sehtation is svLis 3.0, a set theoric based solid modeling kernel [9]. This kernel allows Boolean operations to be performed on solids represented using set theory. Nevertheless, the problem at hand and the advances in modern manufacturing processes are now raising new interest in this representation:II Ability to have multiple membership for a point (e.g.

Nevertheless, the problem at hand and the advances in modern manufacturing processes are now raising new interest in this representation:II Ability to have multiple membership for a point (e.g. being part of several materials), II Allow porous models to be created, II Allow one (or more) materials to blend gradually II Allow mechanical/thermal/electrical/... properties to be specified locally4 Volumetric Representations4.1 VoxelsTwo definitions are commonly found for a voxel,the basic unit of a volumetricrepresentation:II A voxel,image is like a pixel, has 0 dimension. Akin to the way that a digital represented by an array of pixels, a volumetric dataset is made of voxels laid out on a regular 3D grid. For every voxel in this dataset, a scalar value quantifies the membership of this voxel to a given reference material.

For every voxel in this dataset, a scalar value quantifies the membership of this voxel to a given reference material. Often, such a representation can be termed 'spatial enumeration.'II A voxel can also be considered to be a cube of small size.There is no strict definition for a voxel, but we will choose to remain consistent with the is a 3D extension of a pixel and thus, has nodefinition of a pixel and say that the voxel dimensionality.4.1.1 Volume RenderingThanks to constant increases in the computing power of desktop computers, Volume Rendering, traditionally reserved to medical imaging, is now finding its way in areas such as failure analysis, computational fluid dynamics and meteorology.

Though volume rendering [10] merely refers to the act of rendering volumetric data, our primary concern is not with this task,· but rather with the underlying representation of the data to be rendered, along with some of the concepts involved in the handling of this data.Initially, volumetric data was gathered from Magnetic Resonance Imaging (MRI) scanners. These scanners gather data from human organs by measuring the energy received from disrupted atoms as they realign after being subject to an out-of-phaseexcitation signal. Since tissues of similar composition exhibit similar responses to this out-of-phase, signal .• consistency is insured in the interpretation of the data. The snapshot of a cross section is developed by using this response as intensity of the pixels constituting the snapshot. a basis for the computation.

The snapshot of a cross section is developed by using this response as intensity of the pixels constituting the snapshot. a basis for the computation. of the light Conceptually, the scanning of the organs is made in all three planes at the same time but a phase shift· in the excitation signal· (for all three planes) generates a phase shift in the restituting signal, which allows cross section images to be isolated. Without computers, a doctors' attention has to191span several cross sections ata time to correctly interpret the 3D dimensional nature of the organs. Stacking the cross sectionsin3D space not .only helped minimize errors of interpretation but also provided doctors and surgeons witha3D image that could be panned, zoomed and rotated to better locate individual details.The.

principles of Volume Rendering are similar to those of raytracing: for every pixel on the view plane, a ray of light is cast orthogonal to the view plane and directed toward the objects to is be rendered. Each elementary element of volume (termed a voxel) intersecting the ray, composited (or combined) with other intersecting voxels to provide the final light intensity of the pixel on the view plane. Since this computation is performed repeatedly for every view generated of the volume, a weight can be attributed for voxels of similar absorption to filter out some tissues while emphasizing other.

Since this computation is performed repeatedly for every view generated of the volume, a weight can be attributed for voxels of similar absorption to filter out some tissues while emphasizing other. A noteworthy recipient of these techniques is the Visible Human [11], where a man's body was entirely digitized through a similar process, resulting in 1800 cross sections of his body (1000 transverse and coronal MR scans, 1878 transverse scans wi corresponding photographs).4.1.2 Voxel Based RepresentationsVolumetric data sets emanating from human organs are extremely dense, and much of the voxels forming the volume are non-empty. Thus there is little incentive to design space efficient data-structures. The main efforts made to improve the storage of volumetric datasets are aimed at improving the speed of volume rendering algorithms by preprocessing the voxels [12], [13] and skipping empty cells [14].

The use of voxels to represent matter raises an interesting issue when it comes to render the outer surfaces of objects. These surfaces exhibit a shape that is often inappropriately captured by voxels. Nevertheless, these surfaces are rendered as if they where described by a conventional polygonal mesh using a normal estimation algorithm and an interpolation kernel [15]. The improvement over 'naIve' volume rendering is certainly significant, but interpolating the surfaces makes it difficult to use this technique for solid modeling where surfaces are often designed with strict tolerances.Additionally, since the voxels are a sampling of physical data, the storage requirements for those sets are severe: a model containing a grid made from 512 arrays of 512x512 voxels uses at the very least 128 M-byte of memory (for 1 bit encoding). The data set is usually very large and thus not easily held in main memory without using some form of virtual memory.

The data set is usually very large and thus not easily held in main memory without using some form of virtual memory. Another application of a voxel based modeling scheme for solids is also presented in [16], where the aim is to see how a voxel-based modeling fits within an RP enabled manufacturing environment.It should be noted that the use of voxels for solid modeling does greatly simplify CSG operations on solids [14]: Boolean operations performed on solids are reduced to plain voxel to voxel operations whose outcomes are extremely simple to compute. For instance, an empty voxel intersecting a full voxel is an empty voxel, an empty voxel 'union' a full voxel is a full voxel.Octrees Samet [17] proposed to use a hierarchical space-partitioning scheme in order to store volumetric data, specifically the Octree.

The principles of an octree are simple: to recursively subdivide a cube into 8 smaller cells 1/8th the size of the original cube, until either the cube is192empty or its content is below a set threshold size. Ayala et al. [18] then proposed .to perform Boolean operations and rigid body motions on octrees. Although the octree encodIng has an interesting potential, is in the termination criteria of the decomposition sequence that the complexity lies. Some example termination criteria, for 3D polyhedra, are:itA single vertex in the cell, • A single edge in the cell, • A single face contained in the cell, •Several faces sharing the same edge in the cell (which mayor may not be in the cell, Fig.3a).In the case of its 2D counterpart, the quadtree, it is demonstrated that regardless of the termination criteria used, the number of cells is proportional to the perimeter of the object to be decomposed and to the resolution of the decomposition ([17], Chapter 1).

Thus, for complex closed hollow 2D polygons, a quadtree decomposition can yield a large number of cells, which restricts its use.3a. Sample Termination Criteria for QuadTree (single edge sharing the same vertex)3b. Cell Tree Decomposition of a Convex PolygonFigure 3a-3b: Sample Termination Criteria for Quadtree & Cell Tree DecompositionA solid modeling application using an octree based representation scheme is presented in [19]. In this application the termination criteria of the decomposition is to have a single face contained within each cell. Once again, a noteworthy benefit from this volume based decomposition is the simplicity of Boolean operations on solids: though these are slightly more complicated than the operations performed on voxels (cells of different sizes may require additional their outcome is easily predicted.

Equally important, rigid body transformations treatment), (rotations, scaling and translations in 3D) are also described without directly applying a transformation matrix to the coordinates of the vertices of the objects and decomposing it (the operation is performed directly on the octree).1934.1.4 The Cell TreeThe Cell Tree [20] is an encoding for general polyhedral pointsetsofarbittarydimensions (bounded or unbounded). This encoding represents a polyhedral bythe algebraic sum of simpler, convex polyhedra. (holes are·' subtrapted').• Each •convex .polyhedral chain is described by the intersection of halfspaces and represented in a vector (Fig. 3b). A halfspace partitions an n dimensional space in halves, and is a hyperplane of dimension n-1. The representation is assumed to be minimal: a halfspace not intersecting any other halfspace (e.g. empty intersection) is removed from the vector.

The representation is assumed to be minimal: a halfspace not intersecting any other halfspace (e.g. empty intersection) is removed from the vector. Also, since the description is minimal, a halfspace is a boundary of the convex polyhedral it is describing. To further normalize this representation, the halfspaces used in the description of a polyhedral are listed in a single location. A convex polyhedral is then represented using a list of 1s, -1 s and Os referring to whether or not this polyhedral is respectively using a given halfspace, its opposite or not using it.Performing Boolean operations on polyhedra is then a slightly more complicated matter than with the previous representation schemes, though it does not necessitate elaborated algorithms. The union of two polyhedra involves merging the two databases of halfspaces and adding more convex polyhedra to the chain.

The union of two polyhedra involves merging the two databases of halfspaces and adding more convex polyhedra to the chain. Subtracting a polyhedral is a similar task, since the description allows unbounded polyhedra, the complementary of the polyhedral to be .subtracted is created by negating the list of vectors it uses, and then added to the convex polyhedra chain.It is interesting to note that this representation offers some features that other representation do not: in the case of 3D polyhedra, there is no need to evaluate the vertices or the edges, strictly speaking, the entire shape is represented using only planes.

However, a very stringent requirement in this representation is that every 3D solids must have a convex decomposition.4.1.5 Geoscience ModelingGeoscience modeling shares some similitude with mechanical CAD, especially in the area of domain (2D or 3D) representation: 2D geographic domains can be represented by way of polygons, and a 3D hill can be represented with a set of 2D elevation maps. The added dimension of attributes such as soil composition, humidity and erosion tends to make this modeling certainly very relevant for the problem at hand. The critical aspects relevant to us in geoscience modeling [21] are those of uncertainty and fuzziness of the boundaries of the objects described.According to [22] there are two main geographic data models used: the exact model and the continuous field model (Figure 4a-4b resp.). The exact model is made of adjoining polygons tagged with a number of attributes representing a region with set properties (such as.

The exact model is made of adjoining polygons tagged with a number of attributes representing a region with set properties (such as. soil composition, instead of attaching the attributes to a pollution... ). Conversely, the continuous field model, topology, assumes that each attribute is a continuous and. smooth varying function over space. In practice, the function is often discretized on a regularly spaced grid and the overlay of all the grids should provide a meaning similar to the one of the exact object model (with the exception of the boundaries). Note that [23] retained an approach resembling the exact field model in his treatment of multiple material solids.1Property #1Property #2Range of Values for Attribute #2Property #3Property #4forProperty #5Model4b. The Continuous Model (discretized),ILl'.......,......."" 4a..4b. Geographic Data ModelsStrictly speaking, heterogeneous solids do not have any uncertainties associated with boundary definition.

The Continuous Model (discretized),ILl'.......,......."" 4a..4b. Geographic Data ModelsStrictly speaking, heterogeneous solids do not have any uncertainties associated with boundary definition. Evidently, the designer of a part knows exactly where a material is and where it isn't. It is hoW for example a soil boundary is modeled and how the I.ILH.l.LL;::;'1.I of a soil (which is often gradual) is treated in terms of fuzziness of the set (the uncertainty, measurement errors, is not relevant here). attains maximal membership in a given attribute, while the boundary is defined as the region attaining between 0 and the maximal membership (the latterinterval being open on both ends). The use of fuzzy logic to model the boundary is proposed [24] and a similar approach, pertaining to a gradual change between two materials is certainly an interesting issue. A more elaborate discussion is presented in [22].

A more elaborate discussion is presented in [22]. The discussion to differentiate a 'Boolean region' from a 'fuzzy region' illustrates why the problem of point membership in a region can not always be a true or false answer in terrain modeling.tosoils, the "core" region is defined as the region5Objects Approach at the University of MichiganInitially, the University of Michigan's efforts were aimed at a theoretical approach to representing heterogeneous solids, [23], [25], [26]. representation of a solid made from n materials was viewed as an application from the geometrical space to the primary material space [23]. In this context, an n-dimensional vector is associated to every triple (x, y, z) within the solid the material composition as a under consideration. equals to combination ofthe primary materials, such that the sum of all the one. Regularized set operations are· then defined on these solids.

equals to combination ofthe primary materials, such that the sum of all the one. Regularized set operations are· then defined on these solids. another work several blending functions between primary materials are extracted from the literature [28]. Note that no detailed representation structure of the data is proposed to daten-dimensional vectorrepresentsIt is only recently [27], with the use of a DMD (Direct Material Deposition) apparatus that entire design process: from design to manufacturing while including the attention has shifted to design homogenization to obtain the ideal microstructure pattern distribution. The material is considered at the microstructure level, which evidently translates into large· amounts of data and many variables to consider during the optimization stage.

The material is considered at the microstructure level, which evidently translates into large· amounts of data and many variables to consider during the optimization stage. representatipn lacks the ability to group together a large number of similar microstructures to reduce the amount of data needed.1956 Conclusionthere is a need ~o combine the features of traditipnalmechanical representations with the feat\.lres of volumetnc representations. There is no perfect. solution in any of the prevalent represe~tati?n schemes.' and effort must target either the development of novel schemes or the combmatlon of aVailable representations. Furthermore, once the problem of material representation is resolved, these samy principles can be applied to other· types of attributes such as electrical conductivity or thermal resistance. This will allow heterogeneous solids to be completely and entirely represented.Several representation schemes have been reviewed. As expected,7 References [1].

This will allow heterogeneous solids to be completely and entirely represented.Several representation schemes have been reviewed. As expected,7 References [1]. Swann T.A, Keicher D.M., and Miller W.D. "Laser Engineered Net Shaping (LENSTM) Technology Commercialization." Rapid Prototyping and Manufacturing '98, Dearborn, MI.[2]. Konig O. (1999). "Application of Genetic Algorithms in the Design of Multimaterial Structures Manufactured in Rapid Prototyping," Diploma Thesis, ETH Zurich, Zurich, Switzerland.[3]. Huang J., and Fadel G.M. (1999). "Heterogeneous Flywheel Modeling & Optimization." Journal ofMaterials. & Design(Special Issue: Rapid Prototyping and Manufacturing).[4]. Jacobs P.P. (1996). Stereolithography and other RP&M Technologies, ASME Press.[5]. Deckard C.R. (1988). "Selective Laser Sintering," Ph.D., University of Texas, Austin, TX.[6]. Hoffmann C. M. (1989). Geometric and Solid Modeling, An Introduction, Morgan Kaufmann Publisher Inc., San Mateo.[7].

Dutta of Designed Materials." ASME Energy Sources Technology Conference, Houston, TX.Ghosh A.K., Kikuchi N., and Mazumder J. "Design, Representation and Fabrication[28]. Bhashyam S., Shin K. Heterogeneous Objects." UM-MEAM-99-08, CADCAM Group, Ann Arbor.and Dutta D., "An Integrated CAD System for Design of197198.

viewed PaperA Hierarchical V-Network Framework for Part Qualification in Metal Additive ManufacturingByeong-Min Roh1,2*, Hui Yang2, Timothy W. Simpson2*, Albert T. Jones3, Paul Witherell31The University of Oklahoma, Norman, OK 73019 2The Pennsylvania State University, University Park, PA 16802 3National Institute of Standards and Technology, Gaithersburg, MD 20899Corresponding Authors: leonroh86@gmail.com, tws8@psu.eduAbstractAdvances in metal additive manufacturing (AM) technologies have enabled greater design freedoms than subtractive manufacturing has afforded. The design freedoms and flexibilities offered by a metal AM system, however, dramatically increase process uncertainties that may also increase part-quality variabilities. Any metal AM part must be tested, validated, and verified to meet quality, safety, and performance requirements. Common qualification methodologies rely on destructive testing, which is neither cost-effective nor efficient.

Common qualification methodologies rely on destructive testing, which is neither cost-effective nor efficient. Much research, which has been conducted on sensing-based, part qualification in AM systems, attempts to maximize the reduction of destructive testing by closely monitoring the fabrication process in real-time. So much “big data” is generated by this increased use of sensors and available measurement sources. However, the use of the data is still hindered by 1) scale & size and 2) uniformity. We propose a hierarchical, V-network framework of quality assurance with the corresponding translation from ex-situ to in-situ part qualifications. This framework offers an innovative, Cyber-Physical System (CPS) that accurately ties models, processes, and measurements together to interpret the sensor data.

This framework offers an innovative, Cyber-Physical System (CPS) that accurately ties models, processes, and measurements together to interpret the sensor data. The framework also supports and guides translation from ex-situ to in-situ quality measurements, thereby providing a systematic structure and focusing on interrelationships between key observations that influence AM part quality. Ultimately the sensor data can support the detection of process anomalies, thus providing a more streamlined and more efficient qualification process than is otherwise possible.1. Introduction1.1 BackgroundMetal additive manufacturing (AM) technologies allow greater design freedom and manufacturability than traditional manufacturing technologies [1-2]. A wide selection of metallic, powder materials is available for different purposes and applications.

A wide selection of metallic, powder materials is available for different purposes and applications. However, parts made from metal AM technologies usually require secondary post-processing or destructive measuring to qualify the part’s performance. Based on different needs and process stages, quality assurance and control methods include many different activities with varying perspectives. Examples include ensuring the correctness of the design model, the mechanical properties, the material structure, the process signatures, and the sensor signals [3-4].Much research has been conducted on quality assurance and control processes in additive manufacturing systems, such as simulation software [5-9], or in-situ failure awareness systems [10-14].

These researches are relevant to 1) computational techniques such as physical simulation, 2) contact coordinate computing [15-16], 3) geometric design [17-18], measurement and analysis [19-20] using optical computing systems, and 4) geometric accuracy analyses using volume-based methods, including computed tomography [21-22]. The computation and simulation techniques are utilized to test products, systems, processes, and concepts. However, current computational simulations for part qualification have1062difficulties in producing optimum results and dealing with uncertainties, specifically for complex geometries.

However, current computational simulations for part qualification have1062difficulties in producing optimum results and dealing with uncertainties, specifically for complex geometries. The three requirements for a simulation are listed below:(1) Reliability: It is essential to simplify the physics equations and iteratively solve them to simulategas and melt-pool flow in metal AM for a specific environment.(2) Prediction: Real physics derivatives from the predicted models because simulation heavily dependson underlaying physical equations(3) Complexity: Computational complexity in the surrogate model requires a powerful and intensivecomputational capability in the simulation.1.2 Our Proposed FrameworkWe propose a V-network framework to address those limitations using what we call “compositional models. It does so by building, testing, validating, verifying the individual models and and composing them as implied in the V-network Framework.

It does so by building, testing, validating, verifying the individual models and and composing them as implied in the V-network Framework. We view this framework as a foundation that is created to accurately reflect the existing, AM cyber and physical infrastructure. The physical infrastructure includes sensors that produce data during the Metal AM process. Gathered sensor data is then relayed to a processing system and applied to the digital network model. That digital network model used to quantify different aspects of the part’s performance. This digital model does so by executing simulations, calculating current performance, and analyzing potential improvements. The outputs from those models can be used as feedback to the actual physical AM.In our view, this virtual infrastructure can also be applied to non-physical objects, processes and systems, what are commonly called digital twins (DTs). These DTs twins “mirror” the actual AM objects, processes and systems.

These DTs twins “mirror” the actual AM objects, processes and systems. But, these DTs can be used as inputs to simulations that can operate on real-time data. Today, creating such DTs is leveraged by both the Internet of Things (IoT) and embedded sensors technologies. They can capture high-level information that can then be integrated into the virtual model. The V-network Framework is, in effect, a virtual environment where ideas can be tested with few limitations. When connected to a real-time, a monitoring platform, the model becomes an integrated, closed-loop system. A system that can be used to inform and drive a real AM-based Cyber-Physical System,Table 1.

A system that can be used to inform and drive a real AM-based Cyber-Physical System,Table 1. Comparison of AM simulation and V-networkAM SimulationV-networkData interactionReal-timeFeedbackSimulation basisPotential parameters input for testingReal-time feedback from a specific product/processScopeNarrow – primarily designWide1.3 Applying the Framework to an AM-based, Cyber-Physical SystemThe convergence of the cyber and physical parts of AM manufacturing is related to a broader concept of technological-manufacturing convergence. The result of this convergence is a “state” with boundaries between design and qualification. Physical components in this “state” extensively interact with cyber components through sensing, monitoring, and computing methods.

Physical components in this “state” extensively interact with cyber components through sensing, monitoring, and computing methods. These methods enable the monitoring1063of the gathered, “big” sensing data and the analysis of that big data to understand the physical behavior of AM manufacturing better.However, in this paper we focus on the more fundamental question, “How do cyber components and AM physical components actually converge? Answering this question requires the integration of computational processes with the physical processes in additive manufacturing. Integration, on the other hand, requires 1) translating cyber components into physical components for data interpretation and 2) physical components into cyber components for monitoring.Conceptually, there are five, major, physical AM components: input, process, structure, property, and performance. There are no generic details on how these components are physically interconnected in metal AM systems.

There are no generic details on how these components are physically interconnected in metal AM systems. The RHS in Figure 1 shows how these individual physical components are, conceptually at least, physically interconnected. The LHS of Figure 1 shows the individual, cyber components and their interconnections. The LHS component side starts with Design Requirements and ends with the Build Parameters.on destru ti e test F uid f ow urfa e te ture e a uationpatteri rostru ture hara teri ationPart eometrPart featureestru ti e test ensin mission P ume eha iorardnessTou hnessBui d en ironmenta parametersFigure 1. Overview of Hierarchical Structure for Part Qualification in Metal AMFigure 1 also shows the different feedforward/feedbackward links between the LHS Virtual components and RHS Physical components.

Overview of Hierarchical Structure for Part Qualification in Metal AMFigure 1 also shows the different feedforward/feedbackward links between the LHS Virtual components and RHS Physical components. For example, the V-network structure shows interactions between Design & Process requirement (left phase) and Measurement & Qualification (right phase) physical phenomenon along with vertical levels such as quality, part, micro/coupon, nano/micro/meso, physics & signature, and parameters & signal using node (variable) and link (connection). Physical interactions, induced by inputs, cause physical behaviors, dynamic phenomena, and process signatures, creating various emissions, and affect the evolution of microstructure, determining the final mechanical property.1.4 The Benefit of Our Proposed FrameworkThe major benefit of our proposed framework is to boost the performance and efficiency of designing, printing, monitoring, testing, and qualifying AM parts.

It does this by creating a Cyber-Physical System1064(CPS) model for metal AM (See the top left part of Figure 2). The P part of that CPS model provides the systematic approach needed to trace measurable, physical AM phenomena and signatures and link them to process monitoring and quality testing. The results from the process monitoring are used to control the AM process parameters. Therefore, the P part of the V-network assumes the existence of a guideline, including test & validation methods, for physically qualifying complex AM parts.The top RHS part of Figure 2, shows a multi-layer architecture for physically qualifying AM parts. AM qualification typically focuses on identifying and monitoring defects in those parts. In metal AM processes, input process parameters play a key role in determining fabricated part quality and performance.

In metal AM processes, input process parameters play a key role in determining fabricated part quality and performance. The reason is there is a causal chain in metal AM is a sequence of events in which any single new input in the chain effects the next link, and every other link, leading up to a final effect (See Figure 2). For example, the change of laser power and scan speed affects process phenomena, determining the microstructure and property formation. However, understanding the cause and effect chain between different levels in metal AM is very complicated because input parameters affect process phenomenon, determining microstructure and the following mechanical property and performance of the part.

Although some recent research progresses in the modeling approaches can help correlate input-process-structure-property-performance, the interconnections between this architecture have yet to be fully established.By implementing this architecture, we are using an assurance model to increase the quality of fabricated AM parts. These results from the quality assurance model quality link the Physical part of the CPS to the Cyber part of the CPS. This quality assurance model is utilized to allow designer and manufacturer to focus on fabrication and part qualification based on their design requirements.

Convergence of Cyber and Physical Components in Additive ManufacturingA convergence network enables the integration of cyber and physical components by node and link connections. These connections combine several components of different natural processes in metal AM to1065control a physical process through sensor monitoring data and feedback mechanisms that adapt themselves to new conditions in real-time.1.5 Paper OutlineThe paper is organized as follows: Sec. 2 reviews a framework for cyber-physical system and uncertainty quantification and part certification in Additive Manufacturing. Next, Sec. 3 presents the formal hierarchy method, ontology, and V-model to illustrate the structure of the V-network. Next, the structured V-network framework is explained to clarify the applications and results in Sec. 4. The article concludes by offering insights and future direction in Sec. 5.2.

Next, the structured V-network framework is explained to clarify the applications and results in Sec. 4. The article concludes by offering insights and future direction in Sec. 5.2. Review of Past WorkMany types of cyber-physical frameworks have been developed - mainly for design and control - by incorporating machine learning techniques in metal additive manufacturing (AM). Different frameworks serve different purposes while each creates duplicate features and generates disparate results [23]. The current frameworks have yet to be fully leveraged to provide comprehensive guidance for the monitoring of, measurement of, detection of, and feedback to the physical additive process. That guidance, which will be based on real-time, sensor and measurement data, must be integrated into a cohesive, cyber-physical framework for monitoring, diagnosing, testing (the Physical part) and meeting QC/QA requirements (the Cyber part).

To build such an integrated framework, it is critical to understand the various AM quality assurance and control components that may interact with each other in the framework. They can interact at the same level or across different levels in the framework. The following sections review past work in a framework for cyber-physical systems, uncertainty quantification, and part certification for metal additive manufacturing.2.1. Frameworks for Cyber-Physical SystemsBuilding a framework for an AM cyber-physical system can help integrate and simulate cyber and physical components with new functionalities in additive manufacturing. Griffor et al. [24] used a broad range of CPS experts to develop the generic CPS framework and applied it to the four application domains: manufacturing, transportation, energy, and healthcare. Their proposed CPS framework integrates and reorganizes components based on cohesive guidance about the identified concepts of “facets” and “aspect”.

Their proposed CPS framework integrates and reorganizes components based on cohesive guidance about the identified concepts of “facets” and “aspect”. Concepts that satisfy the need for a reference CPS description language where standards, and documented applications can be based. The CPS framework targets controlling the process using changing knowledge, integrating research across sectors, and supporting CPS development with new functionalities. Zheng et al. [25] developed a Cyber-Physical Systems (CPS) system architecture that focuses on building and linking digital twins in smart manufacturing. The authors used a tri-model-based approach to concurrently simulate real-world, physical behavior and update the characteristics of the digital twin. These efforts primarily focused on a conceptual, rather than technical, description of the digital twins and their possible applications without the working details.Dillon et al.

These efforts primarily focused on a conceptual, rather than technical, description of the digital twins and their possible applications without the working details.Dillon et al. [26] developed a framework of Web-of-things, which is a set of web standards for the interoperability of different Internet of things (IoT), for cyber-physical systems for integration of digital computation and communication with physical monitoring and control, enabling robust and flexible systems. Nagar et al. [27] investigated the cyber-physical additive manufacturing model by developing a machine learning module to reduce and replace the common defects and reduce the data overload via wireless networks. Lhachemi et al. [28] reviewed augmented reality, cyber-physical systems, and feedback control for additive manufacturing to provide information on feedback control in additive manufacturing and applied feedback theory to enable users to interact better with AM machines. Mahan et al.

Mahan et al. [29] studied simulating cyber-physical systems for design for additive manufacturing to provide a digital testbed for1066future research to characterize the vulnerabilities of manufacturing and develop new methods to effectively identify compromised parts.2.2. Uncertainty Quantification and Part Certification in Additive ManufacturingUncertainty quantification for additive manufacturing is studied by using statistical and computational methods to make predictions. Wang [30] developed an uncertainty quantification method to optimize composition for improving the mechanical property, expanding the choice of alloy in metal additive manufacturing. They analyzed the process-structure-property relationship for 450,000 compositions around the nominal composition of steel alloy and improved the probability of successful AM builds by 44.7%.

They analyzed the process-structure-property relationship for 450,000 compositions around the nominal composition of steel alloy and improved the probability of successful AM builds by 44.7%. Xie [31] studied mechanistic, data-driven prediction using wavelet transforms and convolutional neural networks to predict location-dependent mechanical properties over printed parts based on process-induced thermal history in metal additive manufacturing. The data-driven prediction helps multiresolution analysis and importance analysis to find the dominant, mechanistic features underlying the process. Tapia [32] investigated uncertainty propagation analysis of computational models in metal Additive Manufacturing and proposed a generalized polynomial chaos expansion framework to finite element model to quantify distributions of melt pool dimension.Uncertainty analysis is needed for real-time, process certification, which is leveraged by in-situ sensor measurements and data-driven analytics.

Together, they permit real-time, estimation and characterization of defects. Waller et al. [33] summarized the non-destructive method to meet materials, design, and test requirements during process optimization, real-time process monitoring, finished part qualification, and certification in additive manufacturing. Petrich et al. [34] conducted real-time quality monitoring using multi-modal sensor fusion, facilitating data-driven defect detection in AM. Lu et al. [35] researched compressive sensors for measuring temperature fields; they used them as inputs to heat transfer models and other numerical methods in AM. Spears et al. [36] conducted a review of in-process sensing and monitoring for selective laser melting, specifying significant challenges from multiple input variables that affect part quality. Salama et al.

[36] conducted a review of in-process sensing and monitoring for selective laser melting, specifying significant challenges from multiple input variables that affect part quality. Salama et al. [37] applied the industrial internet of things(IoT) technique to facilitate real-time monitoring and optimize system parameters such as nozzle temperature and filament breakage/runout, leading to reduced maintenance time.Pandiyan et al. [38] studied a semi-supervised approach to differentiate the defect-free regime and anomalies with acoustic signatures. Nassar et al. [39] examined sensor-based defect detection in the directed energy deposition process using optical emission spectroscopy and data acquisition to research the formation of defects during the build process.

However, these research efforts do not directly guide the part certification in different levels of scales(e.g., GD&T, mechanical property, microstructure, phenomenon, signal, and emission) in metal additive manufacturing to support a systematic part certification strategy to qualify part requirements at different levels of scale and monitoring [40]. The main contribution of the current research is conducting the systematic cyber-physical framework for guiding test & validation based on QC/QA requirements at different levels. Specifically, this article considers 1) selection of sensors and data collection based on quality requirements, 2) capability map of a multi-purpose sensor to detect specific defects and then adjust process inputs, 3) guidance of data fusion to integrate Hierarchical data from multiple sensors to produce an essential dataset for ML.3.

MethodologyIn this section, we introduce an overview for the construction of the V-model to build a quality assurance framework with corresponding part requirements. The development and integration of hierarchy1067models are then proposed along with the detailed, hierarchical information in the network, starting from the requirement to design and testing.3.1. V-Model Development3.1.1. V-Model FrameworkV-model framework is an extension of the waterfall model. This framework provides a development process that results in a general verification and validation model for software and systems engineering [41]. The V-model helps visualize the relationships between each phase of the development life cycle and its associated test and validation. The horizontal and vertical axes indicate time or task completeness (left- to-right) and level of hierarchical step (coarsest to fine-grain abstraction), respectively.

The horizontal and vertical axes indicate time or task completeness (left- to-right) and level of hierarchical step (coarsest to fine-grain abstraction), respectively. The process steps move downwards on the LHS and upwards on the RHS to form the systematic V shape.The V-model facilitates a rigorous, structured method, where each phase can be implemented by the detailed connections of the previous phase. The objective of the V-model is to improve the efficiency and effectiveness of the software and system development process. It does this by reflecting the relationship between test activities and development activities. That development process is decomposed into understanding stakeholders’ requirements, performing requirements analysis, designing the initial outline, designing advanced details, and describing required tests in the fundamental development process from left to right on each other counterparts.

Software and system testing are necessary to incorporate testing activity into the entire software and system development life cycle. The V shape proceeds down and then up, from left to right delineating the sequential development and testing activities. The model highlights the existence of different testing levels and depicts how each relates to a different development phase.3.1.2. Hierarchical NetworkA hierarchical network involves decomposing the complex 3D network into discrete planar levels. Each planar level has it own 2D network topology – it can be drawn on the two-dimensional plane (See Figure 3). Each level, or layer, in the hierarchy, provides specific functions that define its role within the overall network. The hierarchical structure helps understand horizontal and vertical interactions among different levels and provides a mathematical representation of connections, optimizing modeled process features to perform specific roles for the domain network layer.

The benefit of dividing a planar network into smaller, more meaningful decomposition is deploying lower-level steps to understand complex problemsLevel 1Level 2Level 3Planar networkHierarchical NetworkFigure 3. Planar Network and Hierarchical Network3.2. Design Hierchaical Structure in V-Model Framework1068The V-network framework is represented as an ontology hierarchy that describes all design requirements and criteria in Figure 4. The ontological hierarchy includes Quality requirement & management (A, A′), Part specification & inspection (B, B′), Mechanical property & testing (C, C′), Microstructural property & characterization (D, D′), Process phenomenon & signature (E, E′) – defined as follows (See Table 2):takeho dersBui dFigure 4.

Definition of Design Requirements & Measurement and QualificationCPSDesign requirements (Cyber requirements)Measurement and Qualification (Physical measurements)Symbol A B C D E F A′ B′ C′ D′ E′ F′Definition Quality requirements Part specification Mechanical property Microstructural property Process phenomenon Process parameter Quality management Part inspection Mechanical testing Microstructural characterization Process signature In-process signals1069Microstructural property & characterization (D, D′): Define acceptable microstructure property (e.g., phase, crystal structure, microstructural orientation) and Determine grain size, size distribution, and phase volume fraction by evaluation of micrographs (e.g., SEM, TEM, EBSD, XRD, XPS).Process phenomenon & signature (E, E′): Determine relevant physical phenomenon (e.g., melting, solidification, heat & mass transfer, vaporization) and capture and analyze expectation observations for measures of effectiveness(e.g., melt pool, scan track, powder bed & printed slice, crack formation).Process parameter & In-process signals (F, F′): Select process parameter to operate AM machine (e.g., laser power, scan speed, layer thickness, spot size) and measure signals during the build process (e.g., radiation, photon level, pressure, pulse, acoustic emission, wavelength, frequency).Stakeholders and Build are defined as follows:Stakeholders (S): individual or group that has an interest and expectation in any decision or outcome of a product, including suppliers, clients, and organizations.Build (G): fabrication activity based on stakeholder requirements and expectations.To translate design requirements to measurement and qualification, the representation of a matrix ofthe V-network framework is essential.SABCDEFGS ABFDCEGFigure 5.

Matrix of V-network frameworkIn Figure 5, the matrix is transformed from network connections of the V-model. The matrix highlights which of the design or process steps have an impact on meeting the requirements. This benefit allows the product and quality users to focus on the critical measurement and qualification, which flow down into a specific requirement level and find a corresponding measurement method for further examination.3.3. Translation in V-model for VerificationThe V-network framework described in the previous sections is utilized in scenarios of 1) Verification of Part Level, 2) Verification of Coupon & Physical Signature Level, and 3) Verification of Signal Level. This section provides the applications of proposed approaches and examines how V-model can help multiple cases. Descriptions of each scenario are presented below:10701Figure 6.

This section provides the applications of proposed approaches and examines how V-model can help multiple cases. Descriptions of each scenario are presented below:10701Figure 6. Scenario of V-network: (1) Verification of Part Level, (2) Verification of Coupon & Physical Signature Level, (3) Verification of Signal LevelScenario 1 (Verification of Part level): To verify the quality of the set of stakeholder (S) requirements in part level, Figure. 6-(1) shows the hierarchical process of verification of part specification using V- network. Quality (A) aspect of requirements can be extracted from stakeholder (S) needs and then decomposed into the part specification (B). The part specification (B) can be tested and validated to determine whether it meets the requirements by inspection (B’).

The part specification (B) can be tested and validated to determine whether it meets the requirements by inspection (B’). The part inspection (B’) results are used to evaluate the performance of quality management (A’) and then make the quality decision by stakeholders (S).Scenario 2 (Verification of coupon and physical signature level): In Figure. 6-(2), the quality requirement (A) and part specification (B) are identified from stakeholder's needs (S), and then required mechanical property (C), microstructural property (D), process phenomenon (E) are examined to meet the quality requirement (A) and specification (B) to be tested and validated by correlating physical measurement. The figures provide guidance for the test and validation step of mechanical property (C) and process phenomenon (E) by mechanical testing (C’) such as tensile test and sensoring signatures (E’) such as melt pool and plume.

Collected measurement data are used to evaluate inspection (B’) and quality (A’) of the part and then make the decision from stakeholders (S).Scenario 3(Verification of signal level): Figure. 6-(3) illustrates how process signals (F’) change by the parameter (F) on the build (G) stage. In order to verify constant process parameter (F), capturing process signal (F’) and emission from build (G) is required.4. ResultsSection 4 demonstrates how our proposed methodology can 1) support and guide in-situ, ex-situ, destructive, and non-destructive, quality measurements, 2) provide a systematic structure, and 3) focus on the interrelationship between those measurements. Measurements that influence the quality of parts produced using metal AM processes. The results of our methodology show how a multi-level, hierarchical, V-network can be used to facilitate accurate part qualification in metal AM.4.1.

The results of our methodology show how a multi-level, hierarchical, V-network can be used to facilitate accurate part qualification in metal AM.4.1. Implementation of V-modelThis section provides a critical implementation and use case, of the V-network framework, which is based on the test, validation, and subsequent qualification, of the complex, metal, AM parts. The result of this use case sequentially presents a guidance for AM part qualification. This guidance creates a two-way bridge. The forward way links design requirements to the actual part requirements. The reverse way links1071relevant measurement procedures and their real-part measurements to determine if those real measurements achieve the desired part quality. This determination is based on an analytical interpretation of the measured quality that is correlated to the desired quality. The major quality focus in this example, is porosity.4.2.

This determination is based on an analytical interpretation of the measured quality that is correlated to the desired quality. The major quality focus in this example, is porosity.4.2. Final Systems of V-modelFigure 7 illustrates a more comprehensive look at our proposed V-model framework for metal AM quality assurance. The framework represents the convergence of AM cyber systems and AM physical systems that comprise metal AM part qualification. The cyber system starts at the top left by first identifying the design requirements from the stakeholders, then understanding process requirements, and finally ending by creating digital twins of next, actual, physical, process parameters. The physical system starts at the bottom by using those process parameters to build, monitor, test, and measure the AM part. Those physical measurements are then used as inputs to cyber components that do the final inspection.

Convergence of Cyber and Physical System on V-model in Metal Additive Manufacturing Part Qualification1072Convergence of the AM cyber systems and the AM physical system is needed to complete the V-model framework. To support this convergence, we have created a Convergence layer that has three components: a Metal AM Ontology, Ontology Network Database, and Ontology Clusters.Cyber SystemsCyber Systems involve computational components in the V-model for metal AM, interacting network of a computational matrix, multi-dimensional process & causality, semantic map, and process map. Requirement of design & process is includedPhysical SystemsPhysical Systems consider physical measurement & qualification by gathering physical data to monitor observations and test part for part qualification. Sensor capability is guided to select associated signatures and monitoring. Signature (e.g., melt-pool) is captured during the printing process and material test(e.g., EBSD) and tensile test.

Sensor capability is guided to select associated signatures and monitoring. Signature (e.g., melt-pool) is captured during the printing process and material test(e.g., EBSD) and tensile test. CT scan and dimensional measurements are conducted for inspection and quality management.ConvergenceThe convergence layer enables components of the V-model to search for relevant information when making decisions. The convergence includes ontology data, analytics clusters for AM process, and hierarchy knowledge information.5. V-Model Case Study of QualificationThis case study starts with a stakeholder’s requirements. These requirements will impact how the target, and final, design variables are affected, and possibly correlated. These design variables will determine which physical measurements will be made to successfully estimate the physical quantities needed to satisfy design requirements at different levels of abstraction.

Implementing the AM V-model translates design requirements into those physical measurements. This implementation leverages the management of and processing of captured data, which can used to identify defects and anomalies. By doing so, the V-model makes the measurements and computations needed for performance monitoring, process control, and part qualification more manageable and more efficient.Figure 8 provides a user’s perspective on capturing the system's behavior as it responds to stakeholder’s quality requirements.

A use case is represented as hierarchical iterative steps in sequence processing of part qualification, beginning with a quality goal of “Towards Fully Solid Part” from stakeholder and ending when that goal is fulfilled by repeated, multiple, processes at distinct levels.1073Plume behavior Spatter signatureScan track temperatureMelt pool geometryMelt pool temperatureMeasuring cross-sectionTensile testingFatigue testingLess poresTensile strengthLow porosity ( 1 )FatigueCrystal structurePhaseMeltingHeat & mass transferCoolingLaser powerSpot sizeScan speedLow porosity Towards Fully Solid PartVisible lightEmissionMechanical waveInfraredElectromagnetic waveer stemsPh si a stemsGrain structurePhase boundaryPhase volume fractionMicrostructural orientationGrain boundaryQuality assuranceCT scanFigure 8.

Case Study of V- ode : “ oid oid spa e in so idified part”The “Towards Fully Solid Part” requirement is decomposed into the Quality, Part, … Parameter & Signal levels of the Cyber System. In this example, Quality is translated into low porosity; Part is translated into less pores; and, Parameter & Signal is translated into laser power, spot size, and scan speed. Each of these six cyber levels is linked directly to a corresponding level of physical phenomena.In Section 4, we illustrated our three proposed ideas. First, we demonstrated our hierarchical structure of each side of our V-model framework. The LHS is a hierarchical cyber system and the RHS is hierarchical, physical system; an AM-based, part-qualification system. Second, we demonstrated the convergence of two individual systems into cyber-physical systems. We used this system for tracing & searching relevant quality information through the three proposed “convergence compoents”.

We used this system for tracing & searching relevant quality information through the three proposed “convergence compoents”. Third, we showed an interactive, multi-layer, by-directional relationship that links the different cyber and physical components. These illustrations, when combined, enabled us to identify, correlate and qualify parts by comparing design requirements to the actual physical measurements for the AM.This “quality assurance” V-model will allow designers and manufacturers to focus on fabrication and part qualification based on their design requirements. Thus, this advantage provides insights into what can be performed with given design criteria and will prioritize measurement capabilities for determining whether AM qualification is fulfilled.6. Conclusion, Discussion & Future WorkIn this work, we proposed a hierarchical, V-network framework for quality assurance and control in metal additive manufacturing.

Conclusion, Discussion & Future WorkIn this work, we proposed a hierarchical, V-network framework for quality assurance and control in metal additive manufacturing. The framework comprises two systems: a cyber-based design system and a physical-based measurement system. Each system has six, linked, hierarchical levels. Each level defines its role in quality assurance, test and validation; and each level exists at a relative, scale size in the framework.1074Each level is linked physically based on the network connections at each level. These six, connected levels and two systems make up our proposed AM, our V-network, cyber-physical system.This research underpinning this framework included guidance in modeling, testing, validating, and verifying methods in cyber-physical systems. The V-network framework guides quality measurement from the requirement for part certification, providing a rigorous hierarchy structure of quality management.

The V-network framework guides quality measurement from the requirement for part certification, providing a rigorous hierarchy structure of quality management. Networked interrelationship describes the connection between requirement, observation, measurement, and quality elements, guiding quantification of quality uncertainty by collecting correct sensing data and processing data fusion. In this paper, the framework was used to measure, to test, and to validate the quality of AM parts: quality in this paper was based on the number of pores in the part.V-network combines the dynamics of the physical processes with those of the sensors, network, machine, and tools, providing description and modeling, simulation, design, process, and analysis techniques for the integrated whole.

Our V-network paves the path for virtual modeling structured physical process, sensor, data, and interconnection to integrate physical components (e.g., sensors, machine, and physical process) and cyber components (e.g., information, data, and computation) for advanced digital twin.Future work includes 1) conducting experiments to implement the V-network framework to expand the types of quality problems such as cracks, under fusion, and over fusion, 2) process modeling to build extensible systems interacting with sensors, data, processes, and quality in additive manufacturing. Our proposed work will coordinate sensing, computing, and controlling physical AM manufacturing systems and infrastructures, connecting them to the AM cyber system, and reflecting the changes in each other.As this work evolves, we believe that our V-network framework can be leveraged to implement data- driven, real-time, prediction and control approach in the AM industry.

An approach that collects sensor data, observes physical behavior based on that data, uses that behavior to drive various measurement methods.AcknowledgmentsThis research was supported by NIST grant 2019-NIST-MSE-01 and a cooperative agreement between the Pennsylvania State University and the National Institute of Standards and Technology (NIST).References[1] Xiao, X. and Joshi, S., 2020. Process planning for five-axis support free additive manufacturing. Additive Manufacturing, 36, p.101569.[2] Xiao, X. and Xiao, H., 2021. Autonomous robotic feature-based freeform fabrication approach. Materials, 15(1), p.247.[3] Roh, B.M., Kumara, S.R., Simpson, T.W., Michaleris, P., Witherell, P. and Assouroko, I., 2016, August. Ontology-based laser and thermal metamodels for metal-based additive manufacturing. In International Design Engineering Technical Conferences and Computers and Information in Engineering Conference (Vol. 50077, p. V01AT02A043).

and Elahinia, M., 2022. Computational modelling of process–structure–property–performance relationships in metal additive manufacturing: a review. International Materials Reviews, 67(1), pp.1-46.[8] Xiao, X. and Joshi, S., 2018. Automatic toolpath generation for heterogeneous objects manufactured by directed energy deposition additive manufacturing process. Journal of Manufacturing Science and Engineering, 140(7), p.071005.[9] Xiao, X., Waddell, C., Hamilton, C. and Xiao, H., 2022. Quality Prediction and Control in Wire Arc Additive Manufacturing via Novel Machine Learning Framework. Micromachines, 13(1), p.137.[10] Chen, L., Yao, X., Xu, P., Moon, S.K. and Bi, G., 2021. Rapid surface defect identification for additive manufacturing with in-situ point cloud processing and machine learning. Virtual and Physical Prototyping, 16(1), pp.50-67.[11] Snow, Z., Diehl, B., Reutzel, E.W. and Nassar, A., 2021.

In International Conference on Remote Engineering and Virtual Instrumentation (pp. 65-73). Springer, Cham.[28] Lhachemi, H., Malik, A. and Shorten, R., 2019. Augmented reality, cyber-physical systems, and feedback control for additive manufacturing: A review. IEEE Access, 7, pp.50119-50135.[29] Mahan, T. and Menold, J., 2020. Simulating cyber-physical systems: Identifying vulnerabilities for design and manufacturing through simulated additive manufacturing environments. Additive manufacturing, 35, p.101232.[30] Wang, X. and Xiong, W., 2020. Uncertainty quantification and composition optimization for alloy additive manufacturing through a CALPHAD-based ICME framework. npj Computational Materials, 6(1), pp.1-11.[31] Xie, X., Bennett, J., Saha, S., Lu, Y., Cao, J., Liu, W.K. and Gan, Z., 2021. Mechanistic data-driven prediction of as-built mechanical properties in metal additive manufacturing.

Hierarchical, Heterogeneous Material CAD Model with Application to Laser SinteringDavid W. Rosen, Namin Jeong, Yan Wang The George W. Woodruff School of Mechanical Engineering Georgia Institute of Technology Atlanta, GA 30332-0405Variations in laser irradiance and local temperatures in laser sintering cause variations in porosity and material microstructure, which can affect mechanical properties of a part. From the design perspective, it is important to model microstructures, mechanical properties, and their relationships so that parts can be analyzed and designed taking into account their as- manufactured condition. In this paper, we propose the modeling of part geometry and microstructure by using a new hierarchical modeling method. A surfacelet transform is introduced to model microstructure. The application of image processing methods enables multi-resolution representations of microstructure.

A surfacelet transform is introduced to model microstructure. The application of image processing methods enables multi-resolution representations of microstructure. Combined with methods from computational materials design, low resolution microstructure representations can be used to compute effective mechanical properties. The models and methods are demonstrated on two examples, a simple fiber-reinforced composite and a laser sintered nylon-12 material.1INTRODUCTIONFor many years, practitioners in the additive manufacturing (AM) industry have complained about the lack of suitable engineering materials. Others note the large variability and unpredictability of mechanical properties in AM processed materials. Both sets of users would benefit from computer-aided design (CAD) tools that integrate material information with geometry. Furthermore, the capability of deriving mechanical properties from the material and geometry information would greatly aid part design and engineering.

Furthermore, the capability of deriving mechanical properties from the material and geometry information would greatly aid part design and engineering. In this paper, we present a new hierarchical, heterogeneous model for CAD that features multi-scale geometry and material information. The model is illustrated with examples of laser sintered (LS) nylon materials.In multiscale CAD systems, ranging from nano (10-9~10-8 m), meso (~10-7 m), micro (10- 6~10-4 m), to macroscales (>10-3 m), the specifications of material compositions and distributions are as important as geometric shapes. At the traditional macroscale in engineering design, geometry and materials can be separately specified. However, at the micro- and meso-scales, material compositions become important in functional realization, such as in composites and functionally graded materials.

However, at the micro- and meso-scales, material compositions become important in functional realization, such as in composites and functionally graded materials. At the nano-scale, geometry and material property become one-to- one correspondent to each other, such as in crystals and metal-organic frameworks. Therefore, we need a concise and precise form to capture material compositions along with geometry so that engineers can design materials in an interactive environment.Even though multi-resolution geometry representations (e.g., subdivision and wavelets) have been developed, the geometric information is only at the same length scale; from this information, representations at different levels of detail (LOD) can be generated.

The major difference between multiscale and multi-resolution modeling is that in a multi-resolution environment, the complete model information is available at the highest resolution, while in multiscale modeling, the highest resolution need not be generated, except when and where it maybe needed. For instance, a multiscale model does not require the entire Boeing 787 airplane model to be available at the atomic scale if one component is required at this scale. Instead, in a multiscale environment, it is necessary to render the model at a specified scale when and where a specified region of a model can be rendered at a desired scale for a specific purpose, such as visualization, finite-element analysis, or manufacturing process planning [i].In this paper, we propose a new multiscale geometric and materials modeling method that uses an implicit representation to efficiently capture internal and boundary information.

This new approach can effectively represent multiscale complex structures as well as material compositions and distributions. It serves as the foundation for modeling structure-property relationships for materials design. We call our modeling approach dual representation or dual- Rep, since it represents both internal and boundary distributions of materials in a unified implicit form. This new model is proposed to support the multiscale zoom process. Fig. 1 illustrates how such a zoom process is important in representing multiscale materials such as bone structures. The external boundaries can be represented as continuous 2D red curves in Fig. 1a. As we zoom in, the zero-width boundary becomes a domain with internal material distributions, as in Fig. 1b. In other words, explicit boundaries with zero width in the previous scale now become continuous distributions of materials with fine-grained features inside the boundary structures themselves.

1b. In other words, explicit boundaries with zero width in the previous scale now become continuous distributions of materials with fine-grained features inside the boundary structures themselves. On the other hand, when zoomed in, it becomes more convenient to model some internal regions where relatively high gradients of distributions exist (i.e., with large contrast between black and white colors in the images) by boundary representation, as in Fig. 1c. With further magnification of continuous regions to the atomic level, the polycrystalline structure emerges, where grain boundaries and defects can be modeled efficiently with boundary representations again, as in Fig. 1d. Therefore, boundary and internal are complementary structures. The purpose of the dual- Rep model is to support such role exchanges of boundary and internal structures at different length scales.

1d. Therefore, boundary and internal are complementary structures. The purpose of the dual- Rep model is to support such role exchanges of boundary and internal structures at different length scales. The enabling core part of the dual-Rep model is a new basis function called surfacelet.(b) Boundary becomes continuous domain(d) Polycrystalline materials with grain boundaries(a) 2D bone structure with boundary (in red) and internal material distributions(c) Boundary appears in continuous domainFig. 1: An illustration of dual-Rep model for multiscale geometry and materials modeling with boundary and internal structures exchanges at difference scales.In the remainder of the paper, we gave a brief background review of relevant work in Section 2. The proposed surfacelet functions are introduced in Section 3. The dual-Rep multiscale heterogeneous model based on surfacelet transforms is described in Section 4.

The proposed surfacelet functions are introduced in Section 3. The dual-Rep multiscale heterogeneous model based on surfacelet transforms is described in Section 4. In Sections 5 and 6, the surfacelet transform is applied to two example microstructures, a fiber-reinforced composite and a nylon-12 laser sintered material, to illustrate its multi-resolution capabilities.2 BACKGROUNDIn this section, we give a brief overview of relevant work to our proposed multiscale materials modeling approach, including heterogeneous modeling in Section 2.1 and multi- resolution geometry representations in Section 2.2.2.1 Heterogeneous Modelingthe modeling of geometry and material composition. Some researchers may include non-manifold topology under the term, which we allow, but do not emphasize. Two general approaches to heterogeneous modeling have been proposed [ii]: discretized and non-discretized approaches.

Two general approaches to heterogeneous modeling have been proposed [ii]: discretized and non-discretized approaches. In the first category, materials and geometry are modeled separately, such as mesh-based and voxel-based methods, where geometry is approximated by volume meshes or voxels [iii], and material distributions are determined by topology optimization or numerical interpolation from control features [iv]. Other researchers applied voxel-based representations that utilized spatial occupancy enumeration of part geometry [v]. General cellular decompositions have also received considerable attention [ vi ], including some with integrated physical property distributions model [ vii ].

General cellular decompositions have also received considerable attention [ vi ], including some with integrated physical property distributions model [ vii ]. Some researchers have generalized the cellular modeling approach to include manufacturing process- related Local Composition Control (LCC) elements [viii].Heterogeneous modeling usually denotesIn non-discretized approaches, some researchers have separated the representation of material compositions and properties from the underlying part geometry [ix]. Others have utilized implicit modeling approaches, which has advantages in that a common mathematical model is used for both geometry and material composition [x]. Shapiro and coworkers have applied the theory of R-functions to show how material composition [xi] can be performed using implicit modeling approaches and extended the model to include distributions of any physically meaningful quantity throughout a part.

A different group proposed a method based on hypertextures [xii] that provides more intuitive user controls, according to the developers.The limitation of all work described here is that the models are not multiscale; at best some can be called multi-resolution, specifically the methods in the first general approach based on geometric decompositions. Even then, the resolutions are limited to the overall part geometry and the particular decomposition into cells, voxels, or mesh elements. In contrast, our proposed approach is both multiscale and multi-resolution, where part representations can be generated at the desired size scale and resolution.2.2 Multi-resolution Geometry RepresentationsThere have been various efforts on multi-resolution geometry representation. The basic ideas are based on functional approximations in the Hilbert space. Choosing appropriate basis functions that have multi-resolution properties is the core research issue.

The basic ideas are based on functional approximations in the Hilbert space. Choosing appropriate basis functions that have multi-resolution properties is the core research issue. The most studied ones are B-splines and wavelets.B-Spline: B-spline is widely used as the basis function in geometric modeling and simulation. Through knot insertion and degree elevation [xiii], B-spline curves and surfaces can be refined with more control points introduced. Reversely, control points can be reduced in approximations.A more general multi-resolution representation is subdivision curves and surfaces [xiv, xv], which are based on the B-spline refinability. That is, the B-spline basis function can be written as a linear combination of the translated and dilated copies of itself.

That is, the B-spline basis function can be written as a linear combination of the translated and dilated copies of itself. Therefore, by recursively refining control points based on some topological and geometric rules, the control polygon or mesh converges to the limit B-spline curve or surface.Wavelets: In the domain of 2D shape representations, wavelets are among the most popular multi-resolution representations. Similar to Fourier analysis, wavelet analysis is to represent and approximate signals (or functions in general) with orthogonal or non-orthogonal basis functions. Both methods can be used to uncover frequency components of signals. Both can be used to represent multi-resolution subspaces and fast algorithms are available for both.

Both methods can be used to uncover frequency components of signals. Both can be used to represent multi-resolution subspaces and fast algorithms are available for both. However, instead of sinusoidal functions in Fourier analysis, the functional space for wavelet analysis is decomposed based on a scaling function with one dimensional variable t for multi-resolution analysis.and a wavelet function( )tϕ( )tψWavelets are self-similar and can be scaled up and down. More specifically, the waveletfunction( ) ψ ,a b t=a1/ 2 ψ−( a−1(t−b))(1)is scaled by a scaling (dilation) factor a and translated by a translation factor b. Although certain forms (e.g. Haar, Daubechies, Morlet, etc.) have been used extensively, is actually general and can be customized for specific problems. Similar operations can be applied to scaling functions.

Haar, Daubechies, Morlet, etc.) have been used extensively, is actually general and can be customized for specific problems. Similar operations can be applied to scaling functions. The most important feature of wavelets is that they are localized in both real (time) and reciprocal (frequency) spaces due to the property of regularity and vanishing moments. With these properties, wavelets together with scaling functions can be used to approximate signals and functions in a compact form (compared to the traditional Fourier analysis), with local features represented more efficiently. Wavelets provide a general approach for multi-resolution functional analysis. In geometric modeling domain, the wavelet transforms were used to describe planar curves with multiple resolutions [xvi].( )tψThe above different multi-resolution modeling approaches capture geometric information in different levels of detail.

However, they do not meet the requirement of multiscale design scenarios, where geometric information as well as material properties should be captured in a way so that users can retrieve either low-granularity global structures or high-granularity local information without rendering the global picture with the greatest details.Our proposed multiscale representation of geometry and materials is based on a hybrid approach of implicit and boundary representations. The general idea is to represent materials distribution in a domain and its boundary with a combination of two types of basis functions. While regular wavelets can be used in internal distributions, a new basis function is introduced to model materials distributions in boundaries, which is called surfacelet. It is a generalization of ridgelet and curvelet, as introduced in the next section.Wedgelet, Curvelets, and Surflet: Wavelets perform well for objects with point singularities in dimension 1.

It is a generalization of ridgelet and curvelet, as introduced in the next section.Wedgelet, Curvelets, and Surflet: Wavelets perform well for objects with point singularities in dimension 1. However they are not effective in dealing with edge discontinuities in dimension 2.Several so-called directional wavelet approaches have been proposed to solve this issue, including wedgelet [xvii], curvelet [xviii], and ridgelet [xix].The wedgelet approach partitions 2D space into squares as building blocks which have line segments. 2D images then can be approximated by a collection of specifically chosen wedgelets. The curvelet function is an extension of the standard wavelet function, which combines the concepts of neural networks, Radon transform, and wavelets. It was developed to compress images containing continuous line or curve segments, where the standard wavelets are not efficient.

It was developed to compress images containing continuous line or curve segments, where the standard wavelets are not efficient. The idea of the curvelet was originally started as ridgelet, where an angular element θ is introduced in the wavelet function asθψ , , a br ( )=( ψ a b x ,cosθ+ysinθ)=a−1/ 2ψ( a1 −(xcosθ+ysinθ−b))(2)so that the parameters include scaling, translation, and orientation. Later, curvelet was constructed in the Fourier space in order to have better controls of resolution in the frequency domain. With the multi-resolution deployment of ridgelets, a parabolic scaling constraint of width ≈ length2 can be introduced so that curvelets are no longer lines with infinite lengths. If wavelets can be thought of as “fat” points with certain widths of local support, curvelets are “fat” needles where orientation information can be captured.In 3D analysis, Ying et al.

If wavelets can be thought of as “fat” points with certain widths of local support, curvelets are “fat” needles where orientation information can be captured.In 3D analysis, Ying et al. [xx] extended 2D curvelet transform to 3D with similar frequency space tilings. Similarly, Lu and Do [xxi] extended contourlets to three dimensions in a discrete space. Chandrasekaran et al. [ xxii ] extended wedgelets to high-dimensional space and approximate functions with polynomial building blocks, called surflets, instead of linear building blocks in wedgelets.In this paper, we propose a new modeling approach to capture multiscale material distributions with a new surfacelet basis function. The proposed surfacelet can be viewed as a 3D generalization of the above directional wavelets.

The proposed surfacelet can be viewed as a 3D generalization of the above directional wavelets. Also different from the purpose of the above efforts, which is to develop efficient approaches to represent and compress edge and surface singularities in 2D or 3D images in the format of pixels or voxels, we are interested in modeling both continuous distributions and surface singularities efficiently in a unified implicit form. Our new surfacelet functions generalize the shape singularities to higher orders such as quadratic surfaces. This allows for more flexibilities of boundary representation. In contrast to the above directional wavelet bases which are constructed in the Fourier domain, we would like to have direct controls in Euclidean space. The surfacelet is described in Section 3.In order to support dual-Rep that allows for multiscale materials representation, a mixture of wavelet and surfacelet basis functions is chosen.

The surfacelet is described in Section 3.In order to support dual-Rep that allows for multiscale materials representation, a mixture of wavelet and surfacelet basis functions is chosen. Wavelet basis functions are used for continuous distributions, whereas surfacelet basis functions are used for surface singularities or boundaries. This approach enables us to capture both internal and boundary domains in a unified form so that seamless zoom operations between scales can be performed.3 SURFACELETWe need to represent both internal and boundary material distributions in multiscale and multi-resolution modeling. The exchange between boundaries and internal structures such as the ones in Fig. 1 cannot be captured efficiently in the forms of regular wavelets or splines. Therefore, a different approach is required. Boundaries can be viewed as surface singularities that are discontinuous in one direction while continuous in other two in 3D space.

Therefore, a different approach is required. Boundaries can be viewed as surface singularities that are discontinuous in one direction while continuous in other two in 3D space. They can berepresented locally as planes, cylinders, ellipsoids, etc. Therefore, we propose surfacelet basis functions as building blocks of boundaries in multiscale modeling.We define a general surfacelet basis function asψa b , ,p( ) r=a−1/2ψ( a1 − ρ b,p( ) r)(3)where function, a is a non-negative scaling factor,is the location in the domain Ω in the Euclidean space,R is a wavelet=r(,x y z ,):ψ →R) bρ , implicitly defines a surface, with the translation factor b and the shape parameter vector m∈p R determining the location and shape of surface singularities, respectively.

As mentioned in Section 2.2, wavelets are efficient in representing one-dimensional point singularities but not two- or three-dimensional ones. In contrast, surfacelets are efficient in two- or three-dimensional singularities but not one- dimensional ones. Therefore, a combination of wavelet and surfacelet basis functions is chosen. Wavelet basis functions are used for continuous internal distributions, while surfacelet basis functions are used for boundary distributions.In this paper, we will focus on the generation and manipulation of surfacelet models of laser sintered microstructure. In contrast, in an earlier paper [xxiii] we discussed the representation of material distributions and microstructures by first specifying materials and properties at a finite number of positions.

In contrast, in an earlier paper [xxiii] we discussed the representation of material distributions and microstructures by first specifying materials and properties at a finite number of positions. Then the distributions within the domain can be interpolated based on selected basis functions.The approach to generating surfacelet models of sintered microstructure can be described as reverse engineering of materials. We use nano- and micro-scale images such as those from atomic force microscopy and electron microscopy to characterize a material’s structure. The boundaries of grains or phases can be detected from the images. In addition, the shading and its gradients in the images provide the information of material distributions.

The boundaries of grains or phases can be detected from the images. In addition, the shading and its gradients in the images provide the information of material distributions. Given that calibration of images is done in the pre-processing step, a mapping from shading in images to properties can be achieved.The Radon transform is an effective method for representing line singularities in 2D and 3D images [xxiv] (i.e., finding straight edges in images). The Radon transform was developed to reconstruct images from CT scans [xxv], which consist of sets of parallel scans where the source and sensor rotate around the target. We generate surfacelet models by first applying the Radon transform, or a Radon alike transform, to the image, to convert line or edge singularities to point singularities.

We generate surfacelet models by first applying the Radon transform, or a Radon alike transform, to the image, to convert line or edge singularities to point singularities. Then, by applying a wavelet transform to the results of the Radon transform, a representation of the image can be produced that is potentially sparse and enables many image processing techniques to be applied. In our case, we want to take advantage of compression methods and the multi-resolution capabilities of wavelets.sameMathematically, the Radon transform in a domain Ω is the integral along the plane (represented as the dash line in 2D), which is perpendicular to a line at angle α, as illustrated in Fig. 4a. The plane and the line intersect at a point which has the radial distance of μ from the origin. Varying the radius μ results in a vector of integral values, Iα(μ) in 2D and Iα,β(μ) in 3D (in 3D, angles α and β. orient the lines along which the integrals are performed).

For a ×pr k( ) l=1 p(∑ ) i j L , ∈k l ,f(i,j)(10)where f(i,j) is the pixel gray-scale value and Lk,l denotes the set of points that make up a line on the lattice Zp { ( { (2 (Zp = {0,1,…,p-1}):})()+ }ikip i Z ∈,mod ,jLl=j=k l .p(11))L=j,lj Z ∈p l ,pIn this formulation, index k takes the place of angle α and l acts as distance μ.(9)The complete surfacelet transform is the result of applying the discrete wavelet transform to each slice (projection) of the Radon transform, which is each rk(l), for l = 0,…,p. As a result, a materials model can be reconstructed from images based on the surfacelet transform defined by Eqns. 7-11. The process of computing surfacelet coefficients from 3D images is illustrated in Fig. 5. From a 2D materials distribution, the finite Radon transform is calculated and arranged in a 3D matrix with k and l as indices.

5. From a 2D materials distribution, the finite Radon transform is calculated and arranged in a 3D matrix with k and l as indices. Then 1D discrete wavelet transforms along the l axis direction are performed for all α’s. The results are surfacelet coefficients for a particular angle.yμRadon Transform1D wavelet transformsαxlmΩkkFig. 5: The general process of surfacelet transforms.5 SIMPLE FIBER-REINFORCED COMPOSITE EXAMPLEThe surfacelet representation and its hierarchical modeling capabilities are illustrated with a simple example of a fiber-reinforced composite material. Figure 6 shows the sample microstructure, with vertical and horizontal fibers spaced 100 μm apart. We assume a typical carbon-epoxy composite material with property values of Em = 2.94 GPa, Ef1 = 234.6 GPa, Ef2 = 13.8 GPa.

The four fibers from Fig. 6 are identified by the four bright spots (highlighted by manipulating the image) in the Radon transform at (k,l) = (0,61), (0,193), (130,61), and (130,193). The index values of k’s here correspond to the angles of 0 and 90 degrees. Applying the biorthogonal spline wavelet (bior1.3 in the Matlab Wavelet Toolbox) to the Radon transform yields the surfacelet representation shown in Fig. 7b.To view a macro-scale section of a part with this simple fiber-reinforced microstructure, the surfacelet representation can be compressed significantly, then reconstructed at a lower resolution. After eliminating 99.95 percent of the surfacelet coefficients from Fig. 7b, the zoomed-out surfacelet representation appears as in Fig. 7c. Fig. 7d shows the result after an inverse wavelet transform is performed, which corresponds to the Radon transform of the compressed microstructure. After performing the inverse finite Radon transform of Fig.

After adjusting the gray-scale values relative to the original microstructure image by scaling the image size, the -1 that effective elastic modulus is 5.26 GPa, which compares favorably to the value of C11 = S11 was computed using Eqn. 12 suitably modified to account for all four fibers, which was 7.4 GPa.50 50100 100l l150 150200 200250 25050 50100 100150 150200 200250 250k ka. Radon transform50 50100 100150 150200 200250 25050 50100 100150 150200 200250 250k kc. compressed (99.95%) surfacelet representation.e. reconstructed compressed image.50100m15020025050100150200250kb. Surfacelet representation (discrete wavelet transform of (a)).50 50100 100150 150200 200250 25050 50100 100150 150200 200250 250k kd. inverse wavelet transform (compare with (b)).Fig. 7.

inverse wavelet transform (compare with (b)).Fig. 7. Fiber-reinforced composite example.6 LASER SINTERED EXAMPLEA section of laser sintered nylon-12 is shown in Figure 8a [xxvii], which will be the sample microstructure used for this example. Note the regions of symmetry in the x > 200 and y > 200 regions (used to increase the image size to be suitable for the finite Radon transform). In LS, it is common for powder particles to be partially melted, not fully melted. Partially melted regions are visible around the small dark clusters, which indicate unmelted particle cores. Regions with clear striations are fully melted areas and tend to have better mechanical properties. The large dark regions are assumed to have significant porosity.Again, we are interested in assessing its elastic modulus at zoomed-out size scales.

The large dark regions are assumed to have significant porosity.Again, we are interested in assessing its elastic modulus at zoomed-out size scales. Similar to the previous example, we perform finite Radon (b) and wavelet (c) transforms, then compress the result of the wavelet transform (surfacelet representation) by 99 percent, as shown in Fig. 8d. The vertical white strips evident in Fig. 8b indicate directions along which a significant number of crystals (or other microstructure features) are oriented. In contrast, black regions indicate directions along which significant porosity occurs.50501001001501502002002502505010015020025050100150200250a. original microstructure image.b. Radon transform of (a).50501001001501502002002502505010015020025050100150200250c. surfacelet representation.d.

original microstructure image.b. Radon transform of (a).50501001001501502002002502505010015020025050100150200250c. surfacelet representation.d. compressed surfacelet representation (99% compression).50501001001501502002002502505010015020025050100150200e. inverse wavelet transform (compare to (b)).f. reconstructed microstructure image.Fig. 8. Laser sintered material example.Taking the inverse discrete wavelet transform of the compressed surfacelet yields Fig. 8e, which is significantly coarser than Fig. 8b, as we would expect. Yet, the compressed representation maintains enough information to reconstruct the microstructure, as seen in Fig. 8f.The effective stiffness (C11) of the compressed surfacelet can be determined in a manner similar to that used for the fiber-reinforced composite of Sec. 5. The gray-scale values in the image in Fig.

5. The gray-scale values in the image in Fig. 8e are used to indicate material types (porous, melted and oriented crystals, and unmelted cores) and their stiffnesses are combined to give the effective stiffness of the region in Fig. 8f. Dark regions in Fig. 8e correspond to porous nylon-12 (assumed C11 = 1200 MPa), bright regions correspond to oriented, melted material (assumed C11 = 2000 MPa, C22 = 1500 MPa), while intermediate to unmelted and partially melted particles (assumed C11 = 1550 MPa). Computing the effective stiffness yields a value of 1600 MPa, which compares favorably with a measured tensile modulus value of approximately 1750 MPa [xxvii].regionscorrespondTo illustrate the integration of a material model with part geometry, the compressed microstructure from Fig. 8f is applied to a section of a tensile test part, as shown in Fig. 9.

8f is applied to a section of a tensile test part, as shown in Fig. 9. The boundaries of a 5x5 mm section, including a curved region, were modeled using implicit geometry such that each boundary separates “inside” from “outside” the part. Boundary models were combined using an implicit intersection operator toFig. 9. 5x5 mm section of tensile bar with LS compressed microstructure.250yield an implicit function of the part section. Regions in the part interior have function values greater than 0, the part boundary has values of 0, and the region outside the part has function values less than 0. Then, the image in Fig. 8f is mapped into the interior of the part. X and Y mirror symmetries were used to tile the part interior in order to give a continuous looking microstructure.7 SUMMARYIn this paper, we introduce a new hierarchical, heterogeneous representation, dual-Rep, to model both the geometry and material microstructure of parts.

The dual-Rep model represents both internal and boundary distributions with a unified implicit form. Surfacelets as new basis functions are proposed to capture boundary information for both the part and the material microstructure (e.g., grain boundaries). A surfacelet model construction method was presented, based on a surfacelet transform, a 2-step process of computing a finite Radon transform, followed by a wavelet transform. Image processing methods were applied to generate lower resolution microstructure models. This particular surfacelet transform was effective in capturing linear microstructure features, such as crystal boundaries. Compression methods were applied to the surfacelet representation to significantly reduce the representation size, while still enabling microstructure reconstruction. Compression ratios of 99 percent and higher were demonstrated on a simple fiber-reinforced composite microstructure, as well as a laser sintered nylon-12 microstructure.

Compression ratios of 99 percent and higher were demonstrated on a simple fiber-reinforced composite microstructure, as well as a laser sintered nylon-12 microstructure. Reasonable estimates of elastic modulus were computed from the compressed representations using rule-of-mixtures methods.REFERENCES[i]Rosen, D.W.: Thoughts on hierarchical modeling methods for complex structures, Computer-Aided Design & Applications, 6(3), 2009, 419-430.[ii] Kou, X.Y.; Tan, S.T.: Heterogeneous object modeling: A review, Computer-Aided Design, 39(4), 2007, 284-301.[iii] Bhashyam, S.; Shin, K.H.; Dutta, D.: An integrated CAD system for design of heterogeneous objects, RapidPrototyping Journal, 6(2), 2000, 119-135.[iv] Koenig, O.; Fadel, G.: Application of genetic algorithms in the design of multi-material structures manufactured in rapid prototyping, Proc. Solid Freeform Fabrication Symposium, pp. 209-217, Austin, TX, August 9-11, 1999.[v] Wu, Z.W.; Soon, S.

GH DENSITY SELECTIVE LASER MELTING OF WASPALOY®K.A. Mumtaz*, N. Hopkinson*, P. Erasenthiran* *Rapid Manufacturing Research Group, Wolfson School of Mechanical and Manufacturing Engineering, Loughborough University, Loughborough LE11 3TU, UK.Reviewed, accepted September 14, 2006AbstractIn this work, high density Waspaloy® specimens were produced using specially assembled laboratory equipment by Selective Laser Melting (SLM). SLM of Waspaloy® powder was performed using a high power pulsed Nd:YAG laser. The laser parameters pulse energy (J), pulse width (ms), repetition rate (Hz) and scan speed (mm/min) were varied. Process parameter optimization was achieved using factorial analysis to investigate the relationship between specific processing parameters and the formation of Waspaloy® specimens. The optimized processing parameters produced Waspaloy® specimens that were 99.3 % dense. The resultant laser melted specimen’s height, width and contact angles were measured.

The optimized processing parameters produced Waspaloy® specimens that were 99.3 % dense. The resultant laser melted specimen’s height, width and contact angles were measured. Specimens were also observed for the occurrence of porosity.1.1.Solid Freeform FabricationSFF is a family of material additive manufacturing processes that enable quick manufacture of physical models from CAD data [1]. The principal advantage of SFF processes is the ability to quickly manufacture parts of virtually any complexity or geometry entirely without the need for tooling. SFF technologies seek to improve upon conventional processing technologies through a reduction in processing steps, reduction in use of materials and reduction in fabrication time and cost.There exists a need to rapidly fabricate functional fully-dense metal parts without the requirement of hard tooling [2].

Several SFF methods have been developed to produce fully functional end use metal parts, these include Electron Beam Melting (EBM), Ultrasonic Consolidation (UC), Selective Laser Sintering (SLS), Direct Metal Laser Sintering (DMLS) Selective Laser Melting (SLM) and Laser Engineered Net Shaping (LENS™) [3].Research efforts have demonstrated the great potential available within laser processing to manufacture metal components with microstructures and mechanical properties equivalent to or superior to conventionally processed materials [4]. The greatest potential for laser based SFF processes may lie within its capability to directly manufacture fully dense metallic components from powder in a single step [5] as well as the manufacturing of specialized functional applications [6]. Schwendner et al. [7] believes laser based SFF is a viable technology for the direct fabrication of metallic components.

Schwendner et al. [7] believes laser based SFF is a viable technology for the direct fabrication of metallic components. This is due to the versatility, accuracy and ease at which high temperatures can be generated by lasers to melt powdered metal [8]. SLM has the potential to fabricate high density fully functionally end use parts in a single step without the requirement of extensive post processing. This is because the process uses a high power laser to completely melt powered metal from a powder bed, enabling high density processing with the requirement of minor surface finishing such as machining and polishing.2201.2.Selective Laser MeltingSLM is a tool-less manufacturing technique that is driven by the need to process near full density metal components with mechanical properties comparable to those of bulk materials [9]. This additive manufacturing technology fabricates parts by selectively melting layers of pre- deposited metal powder using a focused laser beam.

This additive manufacturing technology fabricates parts by selectively melting layers of pre- deposited metal powder using a focused laser beam. The process begins with the deposition of a thin layer of powder. A laser beam with a defined intensity is then scanned across the surface at points that correspond to the cross section of the component. The heat generated by the laser produced a melt pool that causes localized bonding between powder particles [10].As the laser beam moves away from the melt pool, the molten material is solidified forming a fully dense structure. Each layer is melted deeply enough to fuse it to the underlying layer. Another layer of powder material is then spread onto the previously melted layer and the process is repeated. Material particles that are not melted or not fused remain loose and may be removed once the component is completed. In addition this loose powder acts as a supporting agent for the structure.

Material particles that are not melted or not fused remain loose and may be removed once the component is completed. In addition this loose powder acts as a supporting agent for the structure. Figure 1 shows a schematic of this process.LaserOptics Scanning MirrorLaser BeamUn-melted Powder BedPowder DepositorMelted PowderPre-fabricated LayersTableFabricated ComponentFigure 1Selective laser melting processSLM requires no binder or major post processing operation order to achieve final part properties. Single step processing by SLM can produce cost savings realized by the elimination of conventional multi-step thermo-mechanical processing. Design features such as internal cavities or over-hanging features can be made without joined assemblies.

Design features such as internal cavities or over-hanging features can be made without joined assemblies. Hard to process materials such as intermetallics, refractory metals, and high temperature alloys can be processed [8].The potential for using SLM to fabricate full density functional metal components can cut processing time, reduce cost and improve the performance of a component compared to those produced via conventional processing. Due to the SLM continuously improving process parameters (smaller layer thickness, smaller spot size, etc.), Kruth et al. [11] states that the resulting fabricated parts have mechanical properties better than those produced by indirect SLS as well as bulk metal properties. However producing these end use parts in a single step is not an easy process. There are many processing issues and problems that can arise.2211.3.

However producing these end use parts in a single step is not an easy process. There are many processing issues and problems that can arise.2211.3. High density processing using SLMThe majority of functional parts or tools require a high density, as this improves part strength and avoids defects such as crack formation and eventual part failure. A high density part is formed when there is a reduction in porosity/voids within a structure. Porosity usually develops in laser melting processes when trapped gas is not allowed to escape or when powder particles have not fully melted [12]. It has been demonstrated that low porosity can be achieved with higher laser powers, implying that the pore/void level can be reduced with extra heat input [5]. The power level can be increased or energy density can be adjusted by reducing laser feedrate speed or increase repetition rate [13]. Work conducted by Niu et al.

The power level can be increased or energy density can be adjusted by reducing laser feedrate speed or increase repetition rate [13]. Work conducted by Niu et al. [14] shows that decreasing scan feedrate improved part density for laser processed powdered steel.Another important factor that must be considered within SLM to ensure the successful production of multiple layered parts is the wettability of a melt. Wettability affects how successfully fusion between material particles will occur and consequently the level of porosity. Contact angle is the angle made between an individual melted beads of material and the surface to which it has melted (shown in Figure 2). A higher contact angle usually denotes improved wettablility and reduces defects such as porosity and delamination [9]. Another disadvantageous phenomenon arising from a reduced wettability is balling.

Another disadvantageous phenomenon arising from a reduced wettability is balling. It occurs when the molten material does not wet the underlying substrate due to surface tension and can affect surface flow causing fully molten metal to re-solidify as a series of balls [8]. This results in a rough and bead-shaped surface, obstructing a smooth layer deposition and decreasing the density of the produced part.Fused MaterialContact AngleSubstrateFigure 2Cross-section of a fused bead and the contact angle it makeswith substrateThe success of SLM as a rapid prototyping and rapid manufacturing technology will result mainly from its capability to process a variety of materials to a high density without the need of post processing. However parts fabricated directly from a high power laser beam create a variety of problems when processing metal powder due to the high heat input generated by the laser.

However parts fabricated directly from a high power laser beam create a variety of problems when processing metal powder due to the high heat input generated by the laser. These problems include material vaporization, lack of fusion, part porosity, part shrinkage etc. High residual stresses and material shrinkages are created as parts cool [15], shrinkage also makes it difficult to achieve a high geometric accuracy. Material can warp, leading to cracking and delamination. Eliminating residual pores while maintaining surface quality is an obstacle that further complicates matters with the use of high powered lasers. It is due to these factors that the range of commercially used powdered metals for processing via SLM is still limited today [13].

It is due to these factors that the range of commercially used powdered metals for processing via SLM is still limited today [13]. It222is still the subject of major research studies [14] and development in order to gain widespread acceptance within industry [16].Due to the large amount of problems associated with full density processing, it is critical to have knowledge about laser material processing so as to avoid processing difficulties and setbacks. Correct understanding will allow an operator using a SLM system to vary and control the large variety of processing parameters. Achieving a fine balance between the laser processing parameters is something that is fundamental to the production of high density fully functional metallic parts. The operator will then be able to quickly identify any problems within part fabrication and make proactive decisions to improve the process.

The operator will then be able to quickly identify any problems within part fabrication and make proactive decisions to improve the process. This will allow for successful, reliable and repeatable metal fabrication [17].In an attempt to produce high density specimens with the avoidance of processing issues such as part porosity, material vaporization etc., a processing window for each new material needs to be determined experimentally. This investigation used an Nd:YAG pulsed laser to SLM Waspaloy® from a powder bed. Varying process parameters were used to produce single scanned and layered specimens. A factorial analysis methodology was initially used to identify a suitable laser processing region. Part characteristics such as height, width, contact angle and porosity were examined with the aim of developing optimized process parameters for the production of high density Waspaloy® specimens.2.

Experimental Methodology and TestingResearch performed using SLM and other laser based SFF technologies showed that important part features such as build height and width, porosity and microstructure are strongly dependent on the system’s processing parameters including laser power, laser scanning speed, repetition rate etc. [15]. A quantitative understanding of the relationship between independent process parameters and the formation of single beads and layers of Waspaloy® was undertaken. Part density will depend upon the uniformity and repeatability of solidified melt, whether producing single scans, multiple layers or complex parts. Process parameters can be monitored and optimized continuously to obtain the desired part properties.2.1. Material PropertiesWaspaloy® is a nickel-based, age-hardenable superalloy with excellent high-temperature strength and good resistance to corrosion, notably to oxidation. It is used for aerospace and gas turbine engine components.

It is used for aerospace and gas turbine engine components. Current applications include compressors, rotor discs, shafts, spacers, seals, rings and casings, fasteners and other miscellaneous engine hardware, airframe assemblies and missile systems. The Waspaloy® powder used within this investigation had a mean particle size of 63 mμ , its chemical composition is shown in Table 1.Table 1Waspaloy® Nominal Weight Composition (%) [22]Ni Co 54.0 13.5 19.5 4.2Cr Mo Fe 2.0Si Mn Al 1.4 1.0 0.7Ti Cu 0.5 3.02232.2.System SetupA high power 550W Nd:YAG pulsed laser (GSI Lumonics JK701H) operating at a wavelength of 1.06 mμ was used. The laser beam had a spot diameter of 0.8mm. The beam was carried through a fibre optics delivery system and was installed on a 4-axis CNC controlled machine. Powder layers were deposited in one direction by means of a hopper that traversed across the powder bed. Argon gas was used as a shield gas to prevent parts from oxidization.

Powder layers were deposited in one direction by means of a hopper that traversed across the powder bed. Argon gas was used as a shield gas to prevent parts from oxidization. Parts were built on 43mmx30mmx4mm steel substrates.2.2.1.Single Bead TestingProcessing parameters pulse width (ms), pulse energy (J), repetition rate (Hz) and scan speed (mm/min) all hold an impact upon the formation of SLM parts. Initially single beads of Waspaloy® were processed from a powder bed thickness of 0.9mm to gain an understanding about the laser material interactions and how the powdered Waspaloy® was subsequently formed using varied processing parameters. The single scans were 25mm in length and processed within the range of processing parameters shown in Table 2. These laser processing ranges were based upon work conducted by Su et al.

The single scans were 25mm in length and processed within the range of processing parameters shown in Table 2. These laser processing ranges were based upon work conducted by Su et al. [18] when using SLM to process H13 tool steel to a high density.Table 2Single Scan Processing ParametersPulse Width (ms) 2 - 8Pulse Energy (J) 5 - 15Repetition Rate (Hz) 5 - 15Scan Speed (mm/min) 150 - 250The peak power, specific energy density and spot percentage overlap values generated withinthis investigation are defined by equations 1-3.Peak Power (W) =PulseEnerg Jy )( PulseWidth ( ms )(1)Specific Energy Density =AveragePow × smm )/Feedrate(Wer ( Spotsize)(mm)(2)Percentage Overlap =⎛ ⎜⎜ ⎝spotsize(feedrate × ) mm)/ smm repetition(rate(Hz)−1⎞ ×⎟⎟ ⎠100(3)The processing region in which the material would fully fuse, not bond and vaporize at varying specific densities was initially investigated.

The influence these processing parameters had on the response variables bead width, height and contact angle was measured (see Figure 3) using image tool analysis software. The percentage beam spot overlap shown in Figure 3 was varied between 60-85%. Its impact upon the response variables and wettability were also examined.224Bead WidthScan directionContact AngleαMelted Waspaloy BeadBead HeightPercentage Spot OverlapSubstrateFigure 3Cross-sectioned bead measurements and percentage spot overlapFactorial analysis was employed using the software Statgraphics™ for the experimental design and initial analysis of results. Statgraphics™ produced a vast array of information analyzed from experimental results and suggested which processing parameters held the greatest influence upon the formation of single beads of Waspaloy®.

This allowed for an accurate interpretation of results and optimization of pulse width, repetition rate and scan speed values.2.2.2.Layers fabrication testingUsing optimised pulse width, repetition rate and scan speed parameters, layers of Waspaloy® were then fabricated at varying pulse energies of 6-9J with the aim of reducing porosity within the specimen. The layered specimens were 25mm in length, 5mm in width and consisted of three 0.4mm layers. A scanning strategy that had been successfully developed to produce high density steel parts using SLM was used for processing the layered specimen [18]. An example of the scanning and refill strategy used is shown in Figure 4. The strategy helped prevent unnecessary distortion, allowing the part to maintain a level of geometric accuracy and reduced part porosity. Scans 1,2,3 and 4 were initially performed with a 0.5mm spacing followed by a layer refill strategy and laser scans 5,6 and 7 to complete the layer.

Scans 1,2,3 and 4 were initially performed with a 0.5mm spacing followed by a layer refill strategy and laser scans 5,6 and 7 to complete the layer. The effects pulse energies had on the occurrence of porosity and cracks within the specimen were observed using optical scanning electron microscopy (SEM) analysis.ALayer of powder depositedB1st Laser scan1234Laser refill scanLayer refill of powderCD12341526374Figure 4Cross-sectional view of fusion of layers and refill strategies2253. Results and Discussion3.1.Single Scan Results3.1.1.Bonding RegionLaser processing parameters were varied in an attempt to identify bonding regions in which Waspaloy® would fully bond to the substrate. Single beads that were 25mm in length were processed from a 0.9mm powder bed.

Single beads that were 25mm in length were processed from a 0.9mm powder bed. Varying processing parameters caused the powder to fully fuse to the steel substrate, completely vaporize due to a high specific energy or would not fully fuse due to an insufficient specific energy input. Examples of cross sectioned beads that have undergone the affects vaporization, full and partial bonding are shown in Figure 5. During full laser melting the temperature of the exposed powder should exceed its melting temperature. A further increase in temperature would cause the material to vaporize. When this phase transformation occurs, the rapidly moving evaporated powder particles expand and generate a recoil pressure on the molten pool [7]. The recoil pressure associated with plasma formation was observed to move some of the powder away from the interaction zone, causing the laser to penetrate the steel substrate causing little or no bead formation.

Figure 6 shows the bonding region for the material at different pulse width and specific energies, the material is more susceptible to vaporization at shorter pulse widths compared to longer pulse widths. This is a result of the greater heat intensities that can be generated at lower pulse widths. Therefore lower pulse energies should be used when using very short pulse widths in order to avoid the affects of vaporization.As the pulse width increased the specific energy required for complete bonding of Waspaloy® ) regardless of the2increased.

A lack of bonding occurred at specific energies below 5 ( pulse width due to an insufficient heat input./ mmJIt has been shown that heat generated by the laser and its intensity over a defined area had aconsiderable effect upon the fusion of the Waspaloy® powder.(a)(b)(c)Figure 5Cross-sectioned beads, (a) effect of vaporization, (b) full melting and (c) lack offusion226) 2 m m / J (12010080Bonded Not Bonded VaporizedVaporizationy g r e n E c i f i c e p S6040Bonded Region20Not Bonded00123456789Pulse Width (ms)Figure 6Specific Energy (J/mm 2 ) against Pulse Width(ms)3.1.2.Bead Width and HeightFigure 7 shows the effects that repetition rate, scan speed, pulse width and pulse energy hadon the formation of bead width and heights of fully bonded specimens.An increase in repetition rate caused the respective bead widths and heights to increase.

This is because the total amount of energy delivered to the interaction site is increased, this directly effects the melt pool as the surface tension draws local material into the melt.An increased scan speed reduces the energy input per unit length imparted to the material. This drop in energy causes less powder melting to occur due to a decrease in laser material interaction time and subsequently reduces the bead width. Parts produced at quicker scan speeds will encounter a reduction in incident energy that could lead to higher porosity within a part. However scan speeds that are too slow may lead to vaporization of the material as a longer laser material interaction time occurs.At higher pulse widths the duration the material is exposed to the laser is increased. The longer the melt pool is maintained at an elevated temperature the more powder that will be consolidated into the melt pool, producing larger bead widths and heights.

The longer the melt pool is maintained at an elevated temperature the more powder that will be consolidated into the melt pool, producing larger bead widths and heights. This increase in pulse width does not automatically imply a reduction in part porosity or a successfull bonding to an underlying layer due to a decrease in peak power and material penetration depth.Increasing pulse energies between 5-15J produced larger bead widths and smaller bead heights due to an increase in heat intensity, aiding the full melting of powder. The overall volume of the powder available does not change considerably, which allows the bead width to spread out with a reduction in height.227(a)(b)(c)(d)Figure 7Single bead, processing parameter against bead width/height3.1.3.Bead Contact angleFigure 8 shows the effects that varying processing parameters have on the formation of bead contact angles.

An adjustment in repetition rate did not cause the contact angle made between melt and substrate to vary a significant amount. An increase in scan speed produces smaller, flatter beads due to the surface tension, which increases the contact angle. An increase in pulse width produces a longer laser material interaction time, this increases the amount of material drawn into the melt pool. However the heat intensity would subsequently reduce causing the contact angle to also reduce. This is because there is not enough heat input to fully melt the material and so the molten component does not have enough energy for the capillary forces to spread the material out. An increase in pulse energy allowed for an increase in contact angle. The higher heat input caused the material to undergo phase changes and spread out producing flatter beads with higher contact angles.An increase in repetition rate, scan speed and pulse energy increased contact angle.

Pulse energy held the greatest effect upon contact angle increase and therefore held the greatest influence on the wettability of the melt.228(a)(b)(c)(d)Figure 8Single bead processing, processing parameter against contact angle3.1.4.Spot overlapsThe relationships between bead width, height and contact angle with varying spot overlap are shown in Figure 9(a) shows that the bead width and height varied as the spot overlap was increased. The bead width did not vary a considerable amount between 60-85% overlap. At 70- 80% overlap, a trend in increasing and decreasing bead height was observed. Bead height is at it’s lowest at 65% overlap and it’s highest at 75%. There is a similar trend between the increase and decrease in bead width and height at specific overlaps. Figure 9(b) shows that the contact angle steadily drops and then rises between 70-80%.

There is a similar trend between the increase and decrease in bead width and height at specific overlaps. Figure 9(b) shows that the contact angle steadily drops and then rises between 70-80%. The trend shown in this graph is a reflected image of the trends displayed for both bead width and height values in Figure 9(a). As bead width and height increase, the contact angle reduces. As the bead width and height decrease the contact angle increases. The bead contact angle is highest (103 ° ) at an overlap of 65%, achieved at a scan speed of 168mm/min and repetition rate of 10Hz.(a)(b)Figure 9Single bead, percentage overlap against (a) width/height and (b) contact angle229The single bead investigation allowed for certain process parameters to be optimised. Pulse width 5ms, repetition rate 10Hz and scan speed 168mm/min were found to be the best suited parameters for the production of consistent well formed beads with a high contact angle.

Pulse width 5ms, repetition rate 10Hz and scan speed 168mm/min were found to be the best suited parameters for the production of consistent well formed beads with a high contact angle. The effects of varied pulse energies were further investigated and optimised with the production of multiple layers of Waspaloy®.3.2.Layers of Waspaloy®Layered specimens that were 25mm in length, 5mm in width and consisted of 3 layers(0.4mm layer thickness), this specimen is shown in Figure 10.25mmFigure 10 Layered specimenFigure 11(a) shows the cross-sections of the fabricated specimens and the views under SEM analysis when processed with varying pulse energies. There appears to be a large amount of dimensional distortion throughout the fabricated specimens. This could be a result of the layer thickness, a high pulse width causing too much material to be drawn into melt, scan spacing, high peak power vaporizing material etc.

This could be a result of the layer thickness, a high pulse width causing too much material to be drawn into melt, scan spacing, high peak power vaporizing material etc. However the geometric accuracy of the part is outside the scope of this investigation.The darker areas or spots within specimens shown in Figure 11(a) represent the porous parts of the structure. Two types of porosity can exist within laser processed powders. A porosity due to a lack of fusion or a gas porosity that is caused due to an entrapment of gas or release of gas present within the powder particles. The Waspaloy® specimen’s porosity mainly occurred along layer boundaries and the boundary between single scans within the same layer. Figure 11(b) shows that the specimen processed at 6J pulse energy contained an average porosity of 2.4%. The specimens porosity reduced as the pulse energy was increased.

Figure 11(b) shows that the specimen processed at 6J pulse energy contained an average porosity of 2.4%. The specimens porosity reduced as the pulse energy was increased. This is because the extra heat input improved the melting of the material within the melt pool and reduced part porosity. Parts were produced to an average density of 99.3% when a pulse energy of 9J was used. An increase in pulse energy results in a higher attained temperature of melt and promotes a better interlayer connection. The scanning laser beam delivered a correct amount of energy to each layer to sufficiently re-melt the previous layer without over melting the current layer. Pulse energies above 9J were tested, however the pulse energies generated high peak powers causing large amounts of material vaporization.

Pulse energies above 9J were tested, however the pulse energies generated high peak powers causing large amounts of material vaporization. The plasma plume generated caused the protective glass lens on the Nd:YAG laser to crack.2306J - 1200W7J - 1400W8J - 1600W9J - 1800W(a)(b)Figure 11(a) Layered specimen cross-sectioned under SEM analysis (X2.5) and(b) specimen’s porosity(%) against pulse energy(J)4. ConclusionThis paper has examined the effects of laser process parameters on the formation of Waspaloy® single bead width, height and contact angles using SLM. The results have shown interesting trends in the formation of specimens with variable pulse width, pulse energy, repetition rate and scan speed.It was shown that higher levels of porosity were produced at lower pulse energies and the formation of these pores predominantly exists around layer boundaries. A pulse energy of 9J was sufficient to produce a part that contained an average porosity of 0.7%.

A pulse energy of 9J was sufficient to produce a part that contained an average porosity of 0.7%. The porosity conclusions may be either overestimated or underestimated since the porosity analysis was based on an area survey. However, it could provide qualitative results about the relationship between laser processing parameters and the extent of porosity within a structure.There is still a lot of work that can be performed to develop and improve part properties. Part geometry, micro structural properties and particularly surface finish can be further investigated and optimised in an attempt to produce end use metallic parts using SLM.2315. References 1.Hague, R., Mansour, S., and Saleh, N., Material and design considerations for Rapid Manufacturing. International Journal of Production Research, 2004. 42(22): p. 4691. Yevko, V., Park, C.B., Zak, G., Coyle, T.W., and Benhabib, B., Cladding formation in laser-beam fusion of metal powder. Rapid Prototyping Journal, 1998. 4(4): p.

GH DENSITY TI6AL4V VIA SLM PROCESSING: MICROSTRUCTURE AND MECHANICAL PROPERTIES∗ Chang-Jing Kong, Christopher J. Tuck, Ian A. Ashcroft, Ricky D. Wildman and Richard HagueAdditive Manufacturing Research Group, Wolfson School of Mechanical and Manufacturing Engineering, Loughborough University, LE11 3TUAbstractThis paper investigates a density improvement method for Ti6Al4V alloy processed by the selective laser melting method. A modified inert gas inlet baffle has been employed to develop improved mechanical properties for these materials. Comparisons of the top surface and cross-section porosities of solid blocks processed by the original and modified gas inlet baffles indicate that the modified baffle greatly increases the properties of the processing blocks. Results showed that the porosity of the Ti6Al4V alloy was lower than 0.1% by area. The microstructure of the SLM Ti6Al4V alloy exhibited martenstic α' phase.

Results showed that the porosity of the Ti6Al4V alloy was lower than 0.1% by area. The microstructure of the SLM Ti6Al4V alloy exhibited martenstic α' phase. The UTS tensile strength was 920-960MPa and the elongation at the fracture was 3-5%. The fracture surfaces of the tensile samples demonstrated a mixture of ductile and brittle fracture.IntroductionTitanium alloys are extensively used in the aerospace and aeronautical [1], energy [2] and biomedical [3] industries due to its properties, such as high specific strength at elevated temperature, outstanding fracture resistance, excellent fatigue behavior, exceptional corrosion resistance and superior bio-compatility. However, the complexity of the extraction process, difficulty of melting and problems during fabrication and machining of titanium alloys make them expensive compared to most other metals, which limits their application [4]. At present, most titanium parts are manufactured by conventional machining methods.

At present, most titanium parts are manufactured by conventional machining methods. All types of machining operations, such as turning, milling, drilling, reaming, tapping, sawing, and grinding, are utilized, which results in significant energy consumption and material wastage. In recent year, selective laser melting (SLM) processes have been developed to build components [5] with the advantage of being cost effective for producing one off parts with complex shape and structures [6]. Titanium alloys are some of the metal materials under investigation for SLM production [7-11]. As with many other SLM materials, the qualities of the components processed by the SLM methods produced a major concern as compared to those processed by hot work plus post heat treatment [7,12,13]. High-density SLM Ti6Al4V alloys cannot always be achieved due to the processing characteristics of the additive SLM process.

High-density SLM Ti6Al4V alloys cannot always be achieved due to the processing characteristics of the additive SLM process. Good bonding of the melted layers and lines is needed the issues to obtain high density within parts. Hence, it is necessary to understand the processing effects on the parts produced. Materials processed by the SLM process undergo a rapid cooling process. The microstructure of the Ti alloy processed by SLM will be different from those obtained with conventional manufacturing methods. This will also affect the mechanical properties of the components. It is therefore necessary to understand the microstructure and mechanical properties of titanium alloys processed by SLM. In this paper, the authors describe a method to improve the density of Ti6Al4V alloy components processed by SLM 250 system using a modified inert gas buffle.

In this paper, the authors describe a method to improve the density of Ti6Al4V alloy components processed by SLM 250 system using a modified inert gas buffle. The high-density parts processed by the modified SLM have been metallographically and mechanically characterized.Experimental MethodsThe SLM equipment used was a SLM250 system from MTT Ltd, Which is shown schematically in Figure 1. A 200W laser was delivered through an optical train, which includes a variable focus lens,∗ Corresponding author: email, C.Kong@lboro.ac.ukgalvanometer and F-theta lens before reaching the processing platform. The metal powders were spread on to the processing platform during the interval time of laser scanning. The processing chamber was sealed and evacuated before Argon gas was pumped in to provide an inert gas atmosphere.

The processing chamber was sealed and evacuated before Argon gas was pumped in to provide an inert gas atmosphere. Throughout processing, the inert gas was circulated within the processing chamber by using an extractor pump to take gas out of the processing chamber on the left side, and deliver it back though the right side baffle. Oxygen sensors were used to ensure that the O2 was below 900ppm at all times. If the oxygen level was higher than this value, the inert gas valve opens to refill more argon gas into the chamber.Figure 1. Schematic image of the SLM250 system Figure 2. Plasma gas atomized Ti6Al4V powderPlasma atomized Ti6Al4V powder (provided by LPW Ltd., UK) was used to process thin wall and solid blocks using the SLM250 system. Figure 2 shows an SEM image of the powder. Experiments were carried out on both the original inert gas inlet baffle and a modified thin slit gas inlet baffle (Figure 3).

Figure 2 shows an SEM image of the powder. Experiments were carried out on both the original inert gas inlet baffle and a modified thin slit gas inlet baffle (Figure 3). The gap of the thin slit was 300µm (Figure 3b).Thin walls were processed at a ‘real’ laser power of 177W, laser exposure point distance of 50-70µm and laser pulse exposure time of 60-371µs. The thin wall thickness was measured using vernier callipers. Solid blocks with dimensions of 10 × 12 × 10 mm3 were processed at a laser power of 177W, point distance of 60- 65µm, laser pulse exposure time of 60-371µs and hatch spacing of 100-280 µm. The layer thickness was 50µm and the platform background temperature was 150°C.

The layer thickness was 50µm and the platform background temperature was 150°C. The scanning strategy used was ‘meander scanning’, in which the scanning direction rotates 67° clockwise for each layer.Optical microscopy, field emission gun scanning electron microscopy (SEM) and energy dispersive X- ray spectroscopy (EDX) were used to analyze the powder, thin walls, and the top surface and cross sections of the individual solid blocks. X-ray was used to check the metallographic phase of Ti6Al4V alloy’s powder and solid blocks. The samples were cross-sectioned perpendicular to the blocks and the porosity percentages of the blocks were recorded using a low magnification optical microscopy. Recorded images were processed by software to obtain a statistical measurement on porosity percentage of the viewed areas.

Recorded images were processed by software to obtain a statistical measurement on porosity percentage of the viewed areas. Cross-sectioned samples were etched using Kroll’s reagent with a composition of 3ml HF, 6ml HNO3 and 100ml distilled water according to standard ASTM E407-07 and observed by optical microscope. The microhardness was tested with a load of 500 gram and duration time of 15 second.Tensile tests were performed using an Instron 3369 tensile system (Instron Ltd, Bucks, UK). The tensile samples were designed according to standard ASTM E8/E8M-09 with the subsize specimen, i.e. width of 6mm. The thickness of the samples was 3mm. Tensile specimens were processed on the platform in the direction of X,Y and Z direction. Five samples were prepared in each processing direction. The ultimate tensile strength, yield strength and the elongation values are the average values of the test.Figure 3. Inert gas inlet baffle on the processing chamber.

The ultimate tensile strength, yield strength and the elongation values are the average values of the test.Figure 3. Inert gas inlet baffle on the processing chamber. (a) Original baffle, pump at 24Hz; (b) Thin slit baffle, pump at 13Hz.Results and DiscussionThe initial experiments were carried out using the original inert gas inlet baffle (Figure 3a). In order to optimize the processing parameters, thin walls were processed to find the possible parameters to build solid blocks. Figure 4 shows the relationship between thin wall thickness and energy input. The results indicate an approximately proportional relationship between energy input and thin wall thickness. The energy input per unit length is the control factor for the wall thickness. The completed full height thin walls were on those energy inputs higher than 0.45Jmm-1.Figure 4. Relationship between energy input per Figure 5.

The completed full height thin walls were on those energy inputs higher than 0.45Jmm-1.Figure 4. Relationship between energy input per Figure 5. Hatch spacing and the porosity percentage unit length and the thin wall thicknessSolid blocks were processed at laser power of 177W, point distances of 65 and 60µm, exposure times of 200µm and 170µm on various hatch spacings. The energy input per unit length was 0.54 and 0.50 Jmm-1 respectively. The porosity percentage vs. hatch spacing is shown in Figure 5. Minimum porosity percentages were acquired with hatch spacings of 140 or 160µm, depending on the energy input. Samples processed at a laser power of 177W, point distance of 65µm, laser exposure time of 200µs and hatch spacing 160 µm, which has a minimum porosity percentage, was examined both on the top surface and the cross-section. It was found that there still existed large pores on both the top surface (Figure 6a) and the cross-section (Figure 6b) of the block.

It was found that there still existed large pores on both the top surface (Figure 6a) and the cross-section (Figure 6b) of the block. The largest pores reached 500µm. In order to reduce the porosity percentage, more experiments werecarried out at the increased the energy input with various hatch spacings. However, surface morphology observation of the samples (images are not show in this paper) gave no trend of pores to the hatch spacing variation. It was found that the samples on the left side (close to the inert gas extraction side) had better surface quality than those on the right side (close to the inert gas inlet position).(a)(b)PorePore1 mm1 mmFigure 6. Large pores on top surface and in the cross section of the blocksBack-leftBack-rightCentreFront-leftFront-rightFigure 7.

Large pores on top surface and in the cross section of the blocksBack-leftBack-rightCentreFront-leftFront-rightFigure 7. SEM top surface morphology of the position effect with original gas inlet baffleTo verify the position effect on the processing platform, an experiment was designed to put samples on the four corners and the center of the platform. The same processing parameters were used at a laser power of 177W, point distance of 60µm, exposure time of 185µs and hatch spacing of 140µm. The surface SEM morphology (Figure 7) indicates that position affects the surface qualities of these blocks. Samples close to the inert gas extraction (left side of platform) had better surface quality than samples on the centre and samples close to the inert gas inlet side (right side of platform).Figure 8. Top surface morphology of the samples processed using MTT’s original gas inlet baffle(a) 0.45J/mm, (b) 0.55J/mm, (c) 0.60J/mm, (d) 0.70J/mm, (e) 0.80/mm, (f) 1.0J/mmFigure 9.

Top surface morphology of the samples processed using MTT’s original gas inlet baffle(a) 0.45J/mm, (b) 0.55J/mm, (c) 0.60J/mm, (d) 0.70J/mm, (e) 0.80/mm, (f) 1.0J/mmFigure 9. Top surface morphology of the samples processed using the modified thin slit gas inlet baffle (a) 0.45J/mm, (b) 0.55J/mm, (c) 0.60J/mm, (d) 0.70J/mm, (e) 0.80/mm, (f) 1.0J/mmIn order to test the effect of the gas inlet baffles (shown in Figure 3), experiments were processed with the original inert gas inlet baffle and the modified thin slit inert gas inlet baffle. The same series of processing parameters and sample position were used with a laser power of 177W, point distance of 50µm, exposure times of 127-155µs and hatch spacing of 180µm. Samples were put on the centre of the processing platform for both experiments. The surface morphology of the samples processed using the MTT’s original gas inlet baffle and the modified thin slit gas inlet baffle are shown in Figures 8 and 9.

The surface morphology of the samples processed using the MTT’s original gas inlet baffle and the modified thin slit gas inlet baffle are shown in Figures 8 and 9. The results indicate that using the original MTT gas inlet baffle, increasing the energy input can reduce the number of the top pores on the samples (Figure 8), but the pore size is still large even with a high energy input (Figure 8f). Conversely, smooth surfaces can be obtained at almost every set energy input when using the modified thin slit inert gas inlet baffle (Figure 9). A possible reason for the thin slit gas inlet baffle having an effect on increasing the surface quality may be because it provides greater gas velocity across the build envelope. The gas flow with this design can moreeffectively remove fume and fine powders during processing.

The gas flow with this design can moreeffectively remove fume and fine powders during processing. These fume and fine powders floating within the chamber can block the laser energy during processing and when they deposit on the protective lens a further reduction on the input laser energy occurs.Back-leftBack-middleBack-rightMiddle-leftCentreMiddle-rightFront-leftFront-middleFront-rightFigure 10. SEM top surface morphology of the position effect with the modified thin slit gas inlet baffleFigure 11. Energy input per unit length vs. porosity % Figure 12. Cross-section image of block processed with thin slit baffle with thin slit baffleTo test the position effect samples built with modified thin slit gas inlet baffle, an experiment was designed with samples on the centre, four corners and four sides of the platform with the same processing parameters of laser power of 177W, point distance of 50µm, exposure time of 226µm, hatch spacing of 180µm.

The surface morphologies of the samples are shown in Figure 10. No position effect can be observed on the samples processed using the thin slit gas inlet baffle.The porosity percentage of samples processed using the thin slit gas inlet baffle vs. energy input is shown in Figure 11. The porosity percentage decreases with energy input increases. The lowest porosity achieved was 0.1% in area. Cross-section images show that the porosity sizes are much smaller than seen in those samples processed using the original gas inlet baffle. Figure 12 is an example of a sample processed using the thin slit gas inlet baffle. The processing parameters was of laser power of 177W, point distance of 50µm, laser exposure time of 200µs and the hatch spacing of 180µm.

The processing parameters was of laser power of 177W, point distance of 50µm, laser exposure time of 200µs and the hatch spacing of 180µm. The pore (white) sizes in the sample processed using the thin slit gas inlet baffle (Figure 12) are much smaller (less than 100µm) than those seen in sample processed using the original gas inlet baffle (Figure 6b).(a)(b)100µm100µm(c)(d)100µm100µmFigure 13. Microstructure of Ti6AlV SLM solid blocks. (a) 0.45J/mm (b) 0.60J/mm (c) 0.8J/mm (d) 1.0 J/mmFigure 14. XRD pattern of the powder and solid block Figure 15. the energy input vs. microhardness Solid block processed at 0.8Jmm-1Microstructures of the samples processed using the thin slit gas inlet baffle are shown in Figure 13. Elongated grains are observed growing through the deposition layers (50µm). Martensite plates are visible within the grains, which is consistent with results by other authors [7,9].

Elongated grains are observed growing through the deposition layers (50µm). Martensite plates are visible within the grains, which is consistent with results by other authors [7,9]. The pores formed when using a low energy input (Figure 13 a-c) are mainly due to the lack of bonding between layers. X-ray diffraction spectra (Figure 14) indicate the presence of hcp (α′ martensite) phase only, on both the powder and selective laser melted solid block processed with an energy input of 0.8Jmm-1[7,9,14]. The microhardness results from these samples are shown in Figure 15. The microhardness values increase with energy input. The possible reason for this is the reduction of the porosity within the samples (Figure 11 & 13).The tensile testing results are shown in Table 1 and the initial areas of the strain-stress curve for samples prepared in Y and Z directions are shown in Figure 16.

The yield strength (on 0.2% offset) and ultimate tensile strength of the testing results are lower than those by other authors and the elongation at fracture values are similar [8,9]. The fracture morphology in the samples prepared in the X, Y and Z direction are very similar. Figure 17 shows the fracture structure of a sample prepared in the Y-direction. Mixture of ductile fracture (dimple network) and brittle fracture (cleave and shear lip region) was found. The un-melted pores within the tensile testing samples may be one of the reasons for the reduction of the YS and UTS compared with others results [8, 9].Table 1. Tensile test results of the Ti6Al4V alloy in different processing directions Average YS 0.2% offset MpaAverage UTS MpaSLM250Average Elongation %X-direction Y-direction Z-direction800 770938±8.1 967±11 920±5.23.90±0.73 3.45±0.19 4.44±0.94dimple networkUnsealed porecleave and shear lipFigure 16.

Strain-stress curves of for samples prepared Figure 17. Fracture structure of samples built in in Y and Z direction the Y direction sampleConclusionsA modified thin slit gas inlet baffle was seen to solve the problem of the position effect for the samples in different areas of the platform. It also reduced the percentage and size of pores in solid blocks. The lowest porosity percentage of 0.10% was seen in sample with an energy input of 1.0 Jmm-1 and its microhardness was 375 kgmm-2. The microstructure of the solid blocks were hcp α' martensite phase. The UTS was 920-960MPaand elongation was 3-5%. The fracture structure was a mixture of dimple networks, cleave and shear lip regions.References1. R.R. Boyer, An overview on the use of titanium in the aerospace industry, Materials Science and Engineering, A213 (1996) pp103-1142. R.W. Schutz, H.B.

gh Frequency Ultrasonic Non Destructive Evaluation of Additively Manufactured ComponentsN.V.Karthik*, Hengfeng Gu*, Deepankar Pal *, Thomas Starr†, Brent Stucker*Department of Industrial Engineering, †Department of Chemical Engineering, J. B. Speed School of Engineering, University of Louisville, Louisville, KY 40292Abstract Ultrasonic testing of additively manufactured components is useful for non-destructive defect analysis such as porosity, void and delamination detection as well as for analysis of material properties such as density, material strength and Young’s modulus. A high frequency ultrasonic system has been set up on a Fabrisonics Ultrasonic Additive Manufacturing (Ultrasonic Consolidation) machine to measure the material properties after each layer. The same system is also used to perform offline tests of parts fabricated by SLM.

The same system is also used to perform offline tests of parts fabricated by SLM. Traditional material analysis carried out using SEM and optical microscopy is used to validate and demonstrate the effectiveness of the non-destructive testing equipment.1. Introduction 1.1 Ultrasonic nondestructive testing (NDT) Ultrasonic NDT uses the ability of high frequency sound energy to easily travel through several materials and thus help to conduct examinations and make measurements that can be used for flaw detection, dimensional measurements and material characterization. Ultrasonic testing can be divided into contact, non-contact immersion, and non-contact in air. All kinds of ultrasonic systems require a medium be it either a viscous fluid in contact, water/alcohol in immersion or air in non-contact NDT. Ultrasonic signals are typically of three types, longitudinal, shear and surface waves. In this paper we discuss primarily longitudinal ultrasonic signals.

Ultrasonic signals are typically of three types, longitudinal, shear and surface waves. In this paper we discuss primarily longitudinal ultrasonic signals. A typical non-contact immersion ultrasonic setup is shown in fig 1. An ultrasonic signal of a flaw is explained in fig 2. Typical data presentations common to ultrasonic signals are shown in fig 3. Back wall echoes are explained in fig4.Figure 1: showing a typical NDT setup. The setup used for this paper consists of a 3-12 GHz A/D board, JSR Pulser and receiver, ODIS V3.2 software, 20 MHz immersion transducersFigure 2: shows how cracks can be detected from ultrasonic signals. When there is a discontinuity (such as a crack) in the wave path, part of the energy will be reflected back from the flaw surface. The reflected wave signal is transformed into an electrical signal by the transducer and is displayed on a screen.A-SCANFigure 3: shows the typical data representations of ultrasonic signals.

The reflected wave signal is transformed into an electrical signal by the transducer and is displayed on a screen.A-SCANFigure 3: shows the typical data representations of ultrasonic signals. At every point we have the entire Z information of the part and this single signal is known as an A scan. Rastering across the entire top surface of a component gives us the complete 3D representation of the component and this is represented as the B and C scansBack wall 1 Back wall 2Back wall 3Figure 4: shows three back wall echoes from a sample. When ultrasonic waves travel through a sample, part of the energy is absorbed, part is transmitted and part is reflected back. Multiple reflections from the back surface of the sample are known as back wall echoes and occur once sound travels travels 2*thickness of the sampleFor this project, each of the data collection methods described in Figures 1-4 will be used to extract information from ultrasonic signals.

Correlating these signals to material properties and part quality will be discussed in detail in further sections.1.2 ApplicationsSince additive manufacturing (AM) involves layer by layer manufacturing, this provides a unique opportunity for testing each layer as it is being built, if such an NDT system can be setup. NDT methods such as Infrared thermography have shown some potential for testing the shape and size of a melt pool in laser and electron beam based processes but they are able to measure only visual aspects of material transformations. Ultrasonic testing has been shown to be an excellent and efficient tool for detecting flaws, thickness, grain size, density/porosity, and mechanical properties [1-5]. Ultrasonic testing offers high resolution among the available online NDT techniques. The main drawback of ultrasonic NDT is that it cannot function at temperatures higher than 500K.

Ultrasonic testing offers high resolution among the available online NDT techniques. The main drawback of ultrasonic NDT is that it cannot function at temperatures higher than 500K. This makes Ultrasonic Consolidation (UC) an ideal process choice for ultrasonic NDT as it is a solid state process involving no melting. Once an ultrasonic NDT system has been set up and integrated onto a machine (in this case the Ultrasonic Consolidation system) the goal is to achieve closed-loop control.1.3 Close loop controlControl of a system involves a thorough understanding of the system, the parameters which can be controlled and the sensors which measure the response and give feedback. Control systems are of two types, open loop and a closed loop feedback control. Open loops give an initial value into the system and are based on an assumed state and model of the system. Closed loop control takes the feedback from the performance and changes the input based on the current feedback.

Closed loop control takes the feedback from the performance and changes the input based on the current feedback. In case there is a repetitive action (which is the case in UC -- layer by layer welding), it is also advantageous to use an ILC (Iterative Learning Controller) [6].Closed loop control takes several cycles to stabilize since the values input through the mathematical model are often not very close to the experimental values. Thus we propose to use an open loop feed forward control from DDCP-FEM simulations [7] and a closed loop PIDcontrol (fig 5) with an ILC for each layer for each material. The role of the DDCP-FEM simulations is to give an approximate and close value to the optimized parameters while the Closed Loop PID controller is the main feedback control. The ILC will create a database of information for a given material at every layer.Figure 5: A prototype Control loop system setup involving a feed forward FEM simulation and feedback from ultrasonic NDT2.

Integration of systems 2.1 Mounting the system and online NDTIn order to make effective ultrasonic measurements, it is essential to design an integrated setup which holds the ultrasonic transducer right behind the welding head of the UC machine. This gives us access to the X axis motor of the Fabrisonics machine for following the weld. Motion in the Y-axis is accomplished using a precision linear motor and Z axis focusing is achieved using a stepper motor as shown in fig 6.Figure 6 shows the transducer mounted on the left and the X, Y and Z axes of movementSince ultrasonic waves require a medium (preferably a liquid) it is important to design a structure which ensures that the transducer is immersed at all times during the measurement.

The choice for the contact medium was isopropyl alcohol since it is easy to handle, does not corrode materials of interest, is not very expensive and evaporates after we are done with the measurement, which avoids contamination of the foils before the following layer is added. The design of the immersion structure is shown in figure 7.Figure 7: On the top left and center the complete immersion systems is shown. The top right shows the final design which holds the contact medium only where required. The bottom left shows the entire system and how the X, Y and Z axes work in coordination. The bottom right shows the system physically installed.3.

The bottom left shows the entire system and how the X, Y and Z axes work in coordination. The bottom right shows the system physically installed.3. Materials and Testing MethodologyThere are four different domains in which ultrasonic signals are analyzed, namely a) Time domain: Each signal is studied with respect to time in the X-axis which giveswaveform information from which velocity can be derivedb) Attenuation domain: The attenuation of the back wall echoes is studied giving information about how the signal decays in the sample. This depends on absorption and flaw scatteringc) Frequency domain: Spectral analysis of back wall echoes gives important microstructural informationd) Image domain: For visualization of flaws, B and C scans are used as shown in fig 3.In order to use ultrasonic NDT online, master graphs need to be generated for a given material.

A master graph is a correlation of a measured material property with respect to ultrasonic data for the range in which the material property has been measured. Once the master graph is generated, an ultrasonic measurement is performed and the required material property can be correlated to the signal.3.1 Case studyTo generate master graphs 12 samples of EOS GP1 Stainless Steel (17-4 PH SS) fabricated by an EOS M270 were used as shown in fig 8. Six of these samples have the same energy density and six of them have a varying energy density as shown in table 1Figure 8: showing the EOS setup and the test samples3.1.1 DensityDensity was measured using the Archimedes method. The surface of each coupon was ground using a 180 grit sand paper in order to get rid of rough surface layer which may cause bubble entrapment when measuring in the water. Each coupon was measured three times to obtain a good confidence interval.

The top graph illustrates increasing energy density whereas the lower graph illustrates a constant energy density. Standard deviation in density measurements was 0.004 g/cm3 which indicates good reliable data3.1.2 Grain sizeSEM was used for microstructural observation of as-built coupons at different magnification levels after electrolytic etching. Figure 10 shows an array of micrographs in XY and YZ planes of coupons fabricated using various laser power and scan speed levels (parameters are shown as the red labels on the top left corner of each image). The overlapping, bowl-shaped features on the YZ plane are formed in the melt pool solidification process. Thesemelt pool features, which are created by each laser scan, are parallel to the build direction. Parallel columnar austenitic grains can be observed at higher magnification levels due to their very fine sizes.

Parallel columnar austenitic grains can be observed at higher magnification levels due to their very fine sizes. The columnar grains show strongly preferred orientation since they grow along with the heat transfer directions in both XY and YZ planes.The top right corner SEM image is one of the selected images with 16000X magnification for calculation of average diameter of columnar austenitic grain, where γ grain with circular contour grows perpendicular to the cross section of the XY plane. Austenitic grain diameters under both processing parameters set were calculated, the results are shown in figs 10,11.e n a l P Y Xe n a l P Z YFigure 10: Showing representative micrographs in the XY and YZ planes.Figure 11 shows the austenitic grain size (in nano meters) calculated from SEM images, the standard deviation in grain size ranges from 50-100 nm.

Higher grain sizes averaging over 650 nm are marked in blue for comparison to those averaging around 450-500 nm.3.2 Velocity, Attenuation and Spectral parametersEach of the test cubes shown in fig 12 were scanned in the offline NDT mode using a 20 Mhz delay line immersion transducer with a wavelength of 0.3mm. A 4mm*4mm scan was performed in the center of every cube as shown in fig. Jura has shown that anisotropy plays a major role in propagation of ultrasonic waves [8], thus only one orientation was measured. At each point the entire Z information of the part was stored as a raw waveform and a total of 25*26 pixels of data were collected. These 650 waveforms were then analyzed.

At each point the entire Z information of the part was stored as a raw waveform and a total of 25*26 pixels of data were collected. These 650 waveforms were then analyzed. The thickness of the sample was calculated to an accuracy of 0.1mm using vernier calipers and the velocity was calculated according to the pulse echo overlap method (PEO) from equations 1 & 2 [9]where X is the thickness of the sample and τ0 is the time taken for the waveform of the first back wall echo to repeat itself (measured with an accuracy of 1nS). As shown in fig 12 let B1, B2, B3 be the first three back wall echoes, then the velocity is determined by knowing the thickness of the sample and the time taken to travel twice that thickness. Equation 2 is known as the cross correlation of B1 with B2. A representation via figs 13 shows how the distance between peaks in the cross correlation function help to calculate τ0 accurately.Figure 12 shows a sample cube and the area scanned on the left.

A representation via figs 13 shows how the distance between peaks in the cross correlation function help to calculate τ0 accurately.Figure 12 shows a sample cube and the area scanned on the left. On the right we see a representation of backwall echoes and attenuationFigure 13 shows a typical back wall echo and representative cross correlation with the rest of the signal. We observe peaks at backwall echoes 2,3&4. The time between successive peaks in the backwall echoes gives us the accurate velocity of ultrasound from equation 1Ultrasonic attenuation is the rate at which the signal decays as it travels through the same sample, or it is the rate of decay of back wall echoes as seen in fig 12.

It is calculated from the reduction of amplitude of the ultrasonic impulse and is quantified in terms of the attenuation coefficient α dB/mm given by the equation 3in which x is the sample thickness [mm], A0 is the amplitude of the ﬁrst echo [dB], and A1 is the amplitude of the second echo [dB].Frequency domain representations of the back wall echoes are known to give useful information about grain size. The FFT/Auto power spectrum of the first back wall echo as well as the FWHM of the frequency spectrum have been studied.4. Results 4.2.1 Velocity Master Graph According to the literature [10-14], velocity has been correlated successfully to part density/porosity and also to grain size. Ultrasonic velocity however is more sensitive to change in texture and phase while it cannot accurately detect change in grain size.

Ultrasonic velocity however is more sensitive to change in texture and phase while it cannot accurately detect change in grain size. Density calculations have a very good standard deviation but grain size data has a large variance, hence while looking for grain size correlations we look to determine if an average trend can be observed while for correlations with respect to density we look to obtain quantitative results. It was found that when the density levels are low the ultrasonic velocity is significantly lower when compared to 100% dense parts. According to EOS a density of 7.80 g/cm3 indicates fully dense. For the purpose of this study the highest value of 7.87 g/cm3 is considered fully dense.

According to EOS a density of 7.80 g/cm3 indicates fully dense. For the purpose of this study the highest value of 7.87 g/cm3 is considered fully dense. Velocity when plotted vs grain size does not give any significant correlations, while the correlations with density are shown in fig 14.Figure 14 shows the Velocity vs Density of all 12 samples4.2.2 Attenuation Master GraphAttenuation is a very sensitive parameter and is affected by surface roughness, surface parallelism, grain scattering and grain absorption. It must be noted that since ultrasonic signals are collected at only one location, surface parallelism is very important.

It must be noted that since ultrasonic signals are collected at only one location, surface parallelism is very important. Surface roughness of the samples was in an as polished and etched state using a 3µm diamond polisher used for optical and SEM micrographs.Figure 15 shows the Ultrasonic Attenuation Vs Grain size of all 12 samplesFigure 16 shows the Ultrasonic Attenuation Vs Grain size with highly porous samples removedAll measurements have been made in the Rayleigh region where wavelength is orders of magnitude higher than grain size. Ultrasonic parameters in steels have been extensively studied. Papadakis reported a comprehensive set of data on the attenuation and velocity of both longitudinal and transverse waves in hardened and tempered steels as functions of austenitizing temperature and ultrasonic frequency [11]. Klinman used the ultrasonic attenuation due to grain boundary scattering as proxy for grain size in the well-known Hall-Petch relations [15-16].

Klinman used the ultrasonic attenuation due to grain boundary scattering as proxy for grain size in the well-known Hall-Petch relations [15-16]. Ultrasonic attenuation measurements have been applied by Palanichamy [17] to determine the grain size by measuring the scattering induced attenuation which appears as grain noise; rectifying and averaging such data to achieve the absolute amplitude of back scattered signals is possible. This technique is known as the ultrasonic backscatter method, but it requires accurate modeling of coupling correction and diffraction correction.According to Palanichamy, measurements made on AISI 304 SS show that grain scattering causes 90% of the total attenuation. It is important for us to study and correlate the ultrasonic attenuation specifically to grain scattering and grain absorption. The grain size in our samples is of the order of 400-800 nm. Correlations for such fine grained microstructures have not been reported in the literature.

The grain size in our samples is of the order of 400-800 nm. Correlations for such fine grained microstructures have not been reported in the literature. In figure 15 it can be seen that ultrasonic attenuation has seemingly no good correlation with grain size but it must be noted that attenuation is a function of not only grain scattering and grain absorption, but also a function of scattering by pores. In case of the presence many pores, the attenuation value greatly increases. Hence in figure 16 the samples with high porosity values have been removed. These samples have significantly lower ultrasonic velocities, hence they can be easily detected. On removal of these samples we find a much better correlation between attenuation and grain size. Since the grain size of the samples has a high standard deviation, it is a highly satisfactory result that grain size changes of the order of 300 nm can be detected.

Since the grain size of the samples has a high standard deviation, it is a highly satisfactory result that grain size changes of the order of 300 nm can be detected. It is evident that though there is lot of scatter in the data there is certainly a trend which can be detected; the samples with grain size around 450 nm can be distinguished from those with grain size around 700 nm.4.2.3 Spectral parameters Studies by Palanichamy also showed that spectral parameters such as Full width Half maximum (FWHM) of the first back wall echo are indicators of grain size. Several spectral parameters have been studied and plotted. Most of them are seemingly random and show no good correlation to grain size or part density. The reason for this could also be that in studies by Palanichamy all samples had the same thickness, which facilitated for the differences in the amplitude and frequency spectrum of the first back wall echo not affected by external parameters.5.

Conclusions and Future work Through this exploratory work we have shown that the new experimental apparatus can meaningfully correlate ultrasonic signals to material and microstructural characteristics of parts made using additive manufacturing. Ultrasonic parameters such as attenuation and velocity are very sensitive and small physical changes lead to drastic changes in these parameters. It is important to maintain standard experimental procedures including voltage, gain, pulse repetition frequency and filters applied. Good surface finish gives better results which give more information about the internal structure of the component. Surface parallelism is of utmost importance since even a few degrees of change can deflect important signal information away from the transducer.

Surface parallelism is of utmost importance since even a few degrees of change can deflect important signal information away from the transducer. Considering all of these limitations and the sensitivity of the process it iseasy to understand the large variation in results, but even so it could be shown that we are clearly able to distinguish trends in density by using velocity and trends in grain size by using attenuation.Application of these methods to ultrasonic consolidation faces two major hurdles, one being surface roughness and the other being grain noise. Not only is the surface very rough in UC samples, at the sub surface of every layer we find extremely fine grained microstructure evolving into a coarse grain structure in the bulk foil and back to fine grained structure in the next sub surface. This is bound to give significant grain noise. Ultrasonic back scatter method [18] can be used to analyze these signals.

This is bound to give significant grain noise. Ultrasonic back scatter method [18] can be used to analyze these signals. Ultrasonic scattering as a substitute to grain size in the Hall Petch relation as used by Klinman could be an interesting approach. In either case any information from the sample either in terms of scatter noise or grain noise is useful, since it can give us valuable information on further processing.The study discussed in this paper was done using Longitudinal waves,. Study of shear waves has been shown to have better correlations for attenuation vs grain size as well as velocity vs phase. By combining longitudinal and shear waves, yield strength, fracture toughness and correlations for other material properties such as hardness can be calculated.Currently the data obtained from an ultrasonic test performed in ODIS V2.3 software is exported into MATLAB and then analyzed. Work is in progress to develop a system which is controlled completely online in LABVIEW.

Work is in progress to develop a system which is controlled completely online in LABVIEW. This would give us the complete freedom and capability to control the whole data acquisition process and make it much easier to apply signal processing techniques, intelligent gating and closed loop control.5. Acknowledgements We gratefully acknowledge the help of Joe Vicars in the Rapid Prototyping Center at the University of Louisville for his help developing the experimental test apparatus.The authors would like to gratefully acknowledge the Office of Naval Research which funded this work under award N000141110689.6. References[1] J.Krautkramer, H.Krautkramer Ultrasonic Testing of Materials,Springer- Verlag,Berlin,1990.[2] J.L. Rose, Ultrasonic Waves in Solid Media, Cambridge University Press, NewYork,1999.[3] Vary, A. 'Correlation between ultrasonic and fracture toughness factors in metallic materials' NASA TM-73805, Fracture Mechanics, ASTM, STP 677, ed Smith, C.W.

gh green density ceramic components fabricated by the slurry-based 3DP process, Michael Cima1, Jooho Moon1, Scott Uhland1Jason Grau1 Department of Materials Science and Engineering1 Department of Mechanical Engineerini Massachusetts Institute of Technology, Cambridge, MA 02139, Emanuel Sachs2AbstractThe 3DP process has been modified to directly fabricate high green density parts using submicron powder. The slurry-based 3DP process deposits the powder bed by spraying a dispersed slurry of the component material onto a piston. Alumina, silicon nitride, and lead zirconate titanate components with green densities as high as 67% have been fabricated by the slurry-based 3DP process. Solution phase binder systems have proven to be successful for the new process. Substantially improved surface finish over the conventional dry powder-based 3DP process has been demonstrated. Layer heights less than 50 Jlm can be prepared with this process.

Substantially improved surface finish over the conventional dry powder-based 3DP process has been demonstrated. Layer heights less than 50 Jlm can be prepared with this process. Thus, the stepped surface topography commonly observed in solid free form parts is substantially reduced.IntroductionStructural ceramic components require fine-grained starting powders and high green density in order to achieve near theoretical density after firing. Wet processing techniques such as slip casting and injection molding are commonly used in ceramic forming processes to produce parts that meet Structural ceramic components have been fabricated in the past with 3DP using spray-dried powders. These powders are -50 J.lm agglomerates of submicron powder. Spray-dried powders were easily spread into uniform layers, but printed components had packing densities less than 35%.

These powders are -50 J.lm agglomerates of submicron powder. Spray-dried powders were easily spread into uniform layers, but printed components had packing densities less than 35%. Post-processing was required to achieve sufficient density prior to firing.these criteria.The slurry-based 3DP process has been developed to overcome the difficulties of spreading fine submicron powder and also to enable layers as thin as 10 J.lm to be deposited. This paper discusses the slurry-based 3DP process and initial results with different ceramic materials systems. The process is sufficiently generic to be adapted to new materials systems. Our preliminary experiments have been performed on two types of ceramic materials, high purity alumina (several grades of Ceralox HPA) and sintered silicon nitride compositions (HC Stark MIl, 6 wt. % Ah03, 6 wt. % Y203).

% Ah03, 6 wt. % Y203). Firing conditions were 1650°C for 1 hr in air for the alumina samples and 1755°C for 1 hr under 2 psig N2 for the silicon nitride.The slurry-based 3DP processThe slurry-based 3DP process was developed to fabricate high green density components from submicron powders using wet processing techniques. Figure I shows a schematic of the slurry-based 3DP process. The primary difference from the standard 3DP process371is the manner in which the powder bed is deposited. The powder bed in the slurry-based 3DP process is deposited by spraying a dispersed slurry over the piston and drying. Typical slurry volume fractions are between 30 and 35 vol. %. Binder is then printed into the powder bed to define the shape of the component, and the process is repeated until the component is completed.

%. Binder is then printed into the powder bed to define the shape of the component, and the process is repeated until the component is completed. The powder bed is cohesive, and the part is retrieved by redispersing the unprinted regions in an ultrasonic bath.X-Y Positioning System==qpF=~===Heating====llJ'==~ Sluny SprayBinder JetSpraying Slurry==QPI=== Printed Object:iDrying Powderbed===qp= Sluny Spray....~Printing Binder ! Heating•Finished PartSpraying SlurryDrying BinderFigure 1. Schematic of the sluny-based 3DP processTable 1. Densities of alumina powder beds sintered at 1650°C Powder Size 0.5 f.lm l.0 f.lm 0.5 f.lm l.0 f.lmFormation Method Slurry-Based 3DP Slurry-Based 3DP Slip Cast Slip CastDensity 98.7% 99.0% 99.0% 99.2%Powder bed microstructureThe microstructure of the powder bed is controlled by several variables including stability of the slurry, powder morphology, and the raster variables during the spraying process.

Microstructure of powder beds was studied using mercury porosimetry and scanning electron microscopy (SEM). The green densities of powder beds was typically greater than 60% theoretical density and as high as 67%. Mercury porosimetry also indicated that the pore size distribution was uniform with a single population of fine pores. The densities and microstructure of Ah03 powder beds prepared by the slurry-based 3DP process compare favorably with those prepared by slip casting. The densities of sintered372alumina powder beds are compared in Table 1. Polished cross-sections of sintered alumina and silicon nitride powder beds are shown in Figures 2 and 3, respectively.(A)(B)Figure 2. Polished cross-sections of samples sintered at 1650°C. Microstructure of slurry-based 3DP parts (A) are similar to powder beds prepared by slip casting (B). (Scale marker bar is 200 ~).(A)(B)Figure 3. Polished section of a sintered silicon nitride part made by 3DP (A).

(Scale marker bar is 200 ~).(A)(B)Figure 3. Polished section of a sintered silicon nitride part made by 3DP (A). The section shown in (B) has been etched to reveal the structure of the silicon nitride grains.Powder beds prepared by the slurry-based 3DP process do not contain lamination defects from the layered build process, but can occasionally contain a small number of pores in the 2 to 3 Jlm range, as shown in Figure 2A. The occurrence of these pores is related to improperly chosen parameters spraying of the powder bed. This dramatically illustrated by comparing surface profiles of sprayed lines separated by varying distance. Figure 4 shows such profiles for silicon nitride suspensions on a very flat slip cast support. The cross sections are perpendicular to the traversal of the spray head. The deposit from each individual line is trapezoidal in cross section and slightly less than approximately 300 Jlm wide at the top and has a base that is 400 to 500 Jlm wide.

The deposit from each individual line is trapezoidal in cross section and slightly less than approximately 300 Jlm wide at the top and has a base that is 400 to 500 Jlm wide. Two sprayed lines separated by less than this distance will overlap to produce a protuberance in the surface as observed in the surface profiles in Figures 4A and B. Larger sprayed line spacing leaves373a small gap as shown in Figure 4C. Also shown in Figure 4 are sections of sintered silicon nitride parts prepared with the corresponding sprayed line spacing. The occurrence of porosity seems to correlate with the production of gaps between individual sprayed lines. These small gaps between lines may not be filled when subsequent layers are deposited.40 '[ 30 ~ J 201 J10oo300600900=[ (\~I II i}J Distance (J.un)Distmce(J.IlI1)100 J.llD.100 J.llD.~ Figure 4.

Figures A,B, and C are surface profiles of two sprayed silicon nitride lines spaced 200, 250, and 300 !JlIl, respectively, apart on a slip cast bottom layer. Separation of sprayed lines by 300 J.llD. or greater leads to a gap between the sprayed lines. Figures D, E, and F show the corresponding silicon nitride powder beds prepared with sprayed line spacing of 200, 250, and 300 Ilm, respectively, after sintering. The large pores in the microstructure coincide with conditions where the gap forms between two sprayed lines.~~Binder selection and interactionsthe slurry-based 3DP process must meet several The binder systems used for requirements. First of all, the binder needs to penetrate the powder bed and have sufficient strength to survive the part retrieval process. Secondly, the binder must be insoluble prior to redispersing the unprinted powder bed. The intrinsic properties of the binder are also critical to its performance in ink-jet printing and binder-powder interaction.

The intrinsic properties of the binder are also critical to its performance in ink-jet printing and binder-powder interaction. Important intrinsic properties include surface tension, viscosity, conductivity, pH and chemical stability. Resolution of the printed part in the 3DP process mainly depends on binder-jet performance and binder-powder bed interaction. Low surface tension binder has high penetration rate and penetration depth, but it spreads more easily, reducing printing resolution as compared to high surface tension binder.374Three types of commercially available binder systems were examined- polymeric particle suspended binder (latex), wax emulsion binder, and homogenous solution phase binder systems. Latex binders do not infiltrate into the powder bed even if the individual polymer particles have smaller particle size than the mean pore size of the power bed. It is believed that the polymer particles tend to coalesce and form a surface film on the powder bed.

It is believed that the polymer particles tend to coalesce and form a surface film on the powder bed. The smallest particle size of polymeric particles examined was 14 nm in the Estek 1200 polyester latex as determined by dynamic light scattering technique. Even this latex forms a surface film. Wax emulsion binder also forms a surface film when it is printed onto the powder bed. However, during drying, the wax surface' film easily melts and penetrates into the powder bed. Low melting point waxes such as carnauba wax and paraffin are preferred since they must melt to infiltrate the powder bed. Solution phase binders easily penetrate into the powder bed as long as the molecular weight of the binder is less than ......5,000. Solution phase binders must be cured at elevated temperature to strengthen the part and to make the binder insoluble in water.Joncryl 52 is a water-based solution phase binder that has been used effectively.

It is a styrene-acrylic copolymer with molecular weight 1800 and viscosity of 2.8 cps for a 10 volume % solution. Binder burn-out from the part is easily achieved during the firing cycle, and binder residue is not significant in the sintered component. Furfuryl alcohol resin has also been used as a solution phase binder. Furfuryl resin can be diluted in acetone to yield a low viscosity, low surface tension binder. These solution phase binders offer the ability to control both surface tension and viscosity. Penetration behavior of three different binder systems is summarized in Figure 5._ .. _ ..-Drying(A)..Latex Binder (Particle Size = 14nm)Pennanant Surface FilmPrinted Powderbed (Mean Pore Size =100 nm)(B)'Wax Binder (Particle Size =56 nm) ..~-~DryingInfiltrated by Melting-(C)..-Solution phase binderDirect penetrationFigure 5.

Schematic picture showing penetration behavior ofthree different binder systems.Another important aspect of the binder system is the wetting behavior difference between the binder printed region and the unprinted region of the powder bed. Lamination defects375and surface roughness are observed when the wetting behavior differs substantially between the binder printed and unprinted regions. This wetting behavior difference depends on properties of binder materials. Aqueous slurries do not wet regions of the powder bed printed with wax emulsions due to the hydrophobic nature of the wax. Alcohol-based or aqueous-alcohol mixed slurries have a much lower contact angle and wet the binder printed regions more uniformly. Addition of a wetting agent into aqueous slurry also improves the wetting characteristics and consequently gives rise to better printed microstructure as shown in Figure 6.

Addition of a wetting agent into aqueous slurry also improves the wetting characteristics and consequently gives rise to better printed microstructure as shown in Figure 6. Triton X-I00 surfactant decreases the contact angle of 35 v/o alumina slurry on paraffin substrate from 98° to 45° while maintaining a well-dispersed suspension state in the slurry.Layer HeightBuild-up Direction. - 45 J..U1lLayer Height -42~m(A)(B)Figure 6. Cross sectional microstructures of sintered powder bed prepared with and without wetting agent in the slurry are shown in (A) and (B), respectively. Lamination defects are observed in (A) due to the wetting behavior of the slurry on binder printed regions.Printed parts and surface finishThe effect of the amount of binder deposited on the microstructure of the printed part was studied in detail by varying binder flow rate, line spacing, print speed, and layer height.

Single line primitives of binder printed into slurry-based 3DP powder beds have a semi circular cross section compared to circular primitives that are observed for the standard 3DP process. The surface finish was found to be improved by printing with smaller layer heights. Figure 7 shows cross sections of alumina parts printed with 10 weight % Joncryl 52 solution. The edge of the part printed with 50 J.lm layers has a smoother surface finish compared to the part printed with 100 J.lm layers. Single line primitives with this binder are on the order of 150 J.lm thick. The surface finish begins to degrade and the edge of the part has a saw-toothed appearance as the layer thickness approaches the thickness of a single primitive.Surface finish is also dependent upon the surface tension and viscosity of the binder system. Furfuryl resin binder has both low viscosity and surface tension (-23 dynes/em) since it is diluted in acetone.

Furfuryl resin binder has both low viscosity and surface tension (-23 dynes/em) since it is diluted in acetone. Parts printed with furfuryl resin binder display the best surface finish among the binder systems. Figure 8 displays a cross section of an alumina part printed with furfuryl resin. The angled surface does not display stair-stepping that is376typical in most SFF processes, but printing resolution is substantially reduced due to significant spreading of the printed binder.(A)(B)Figure 7. Cross sections of alumina parts printed displaying the effect of layer height on the surface finish of the parts. Parts with 50 J.1In layer height (A) has smoother surface finish than parts printed with 100 J.1In layer height (B). Binder concentration is identical for both parts. (Scale marker bar is 500 J.1In)Figure 8. Cross section of alumina part printed with furfuryl resin binder.

Binder concentration is identical for both parts. (Scale marker bar is 500 J.1In)Figure 8. Cross section of alumina part printed with furfuryl resin binder. Angled surfaces do not display stair-stepping.Parts have been printed with three ceramic materials systems to date: alumina, silicon nitride, and lead zirconate titanate (PZT). Two simple silicon nitride parts are shown in Figure 9 and a bisque fired cross section of one such silicon nitride part is shown in Figure 10. The sample remains porous upon bisque firing and does not shrink. The mounting material is observed to creep into the pores near the surface of the bisque fired sample and is the cause of the different contrast near the sample surface. The bisque fired377section is used to estimate the dimensional accuracy of the slurry based 3DP process. This is a fifty layer part which consists of a flat base with middle section of which rises above the base. The middle section is nine lines wide.

This is a fifty layer part which consists of a flat base with middle section of which rises above the base. The middle section is nine lines wide. A Joncryl 52 solution is used as the binder with a 300 f.lm line spacing. The binder flow rate is such that this spacing results in a part that is 3.4 wt. % binder. Also, the spcaing is such that the edge lines of the middle section are printed 2700 f.lm apart. Measurements of the binder spreading in silicon nitride powder beds reveals that individual lines of printed binder are on average 342 f.lm wide. The expected middle section width is, therefore, 2700 f.lm plus 342 f.lm or 3042 f.lm. The measured value is 3110 f.lm with an absolute error of 68 f.lm or 2%. More work is needed to refine the process planning rules to account for such systematic errors.Figure 9. Two simple silicon nitride parts used for dimensional testing.Figure 10. A cross section of a bisque fired silicon nitride part shown in Figure 9.

Two simple silicon nitride parts used for dimensional testing.Figure 10. A cross section of a bisque fired silicon nitride part shown in Figure 9. The material with different contrast near the surface of the sample is the result of mounting material only partially permeating the porous sample.ConclusionsThe slurry-based 3DP process has been used successfully to prepare high green density components with fine-grained ceramics. Improved surface finish is possible by controlling the properties of the binder system and reducing the layer height. The process is generic and can be readily adapted for new materials systems by optimizing the stability of the slurry. Our future work will be to refine the process planning so as to more precisely control the final component dimensions.378.

gh Pressure, Convectively-Enhanced Laser Chemical Vapor Deposition of TitaniumJames Maxwell Ramnath Krishnan Suresh Haridas Institute for Micromanufacturing Louisiana Tech UniversityLaser chemical vapor deposition is a freeform technique that can generate three-dimensional structures from organometallic or metal halide precursors. To obtain enhanced growth rates, a novel high pressure reactor has been constructed where impinging jets of volatile fluids are heated and pyrolyzed to create parts. Argon Ion and Nd:YAG lasers have been used to selectively generate three-dimen sional titanium shapes from titanium tetrachloride, titanium tetrabromide, and tita nium tetraiodide, at pressures up to 3.0 atmospheres. Emission lines characteristic of the reaction rate have been identified which will allow feedback control of the reaction rate. The process is being optimized to obtain high deposition rates, energy efficiency, and desirable material morphologies.

The process is being optimized to obtain high deposition rates, energy efficiency, and desirable material morphologies. A feedback control system is required to generate 3-D structures with dimensional accuracy and predictable deposition rates.I. IntroductionThis paper reports the development of anew process which holds promise for the desktop manufacture of metal and ceramic parts at the centimeter scale or larger. Termed high-pressure, convectively-enhanced, laser chemical vapor deposition (HPCE-LCVD), the process employs convective flows of precursor gases to enhance the growth rate of vapor-phase solid freeform fabrication (SFF). Preliminary results in the growth of titanium from the titanium halides at moderate pressures and flowrates (up to 3 atmospheres) are presented herein.

5. Careful adjustment of the laser power may be required during the transient growth6 of metals to obtain continuous, uniform dimensions--as will be seen in this paper, where the growth of titanium is thermally-self-limiting at low pressures (see Fig. 3).Several authors have explored the possibilities of three-dimensional growth using LCVD, prototyping rods? fibers,8 coils,9 blocks,10 and more complex structures.II Rods have been grown of aluminum,12 alumina,13 gold,14 nickel,I5 tungsten,I6 silicon,1? chromium-oxide,18 iron,19 steel,2° and boron.21 It is noteworthy, however, that the materials with the greatest ther mal conductivities, e.g. aluminum, gold, and tungsten, are highly irregular in shape, whilst497continuous, steady-state rod growth was readily demonstrated with the other materials.

aluminum, gold, and tungsten, are highly irregular in shape, whilst497continuous, steady-state rod growth was readily demonstrated with the other materials. Mate rials with high thermal conductivity also tend to grow more broadly, or in the extreme case, only as hemispheres.22 Thus, controlled growth of highly-conductive metals is a needed goal.One method for increasing the aspect ratio of a metal deposit is to use a pulsed-laser or a chopped beam.23 This method was employed throughout the Ti growth experiments. Another method for obtaining controlled freeformgrowthwas previously demonstrated by the author;24 where spectral emissions generated during an LCVD reaction were measured to determine the volumetric growth rate. This measure was then used to automatically compensate for errors in the growth, adjusting the incident laser power and other process inputs in real-time.

This measure was then used to automatically compensate for errors in the growth, adjusting the incident laser power and other process inputs in real-time. This tech nique will be required for controlled growth of titanium, whose bulk thermal conductivity is 21.9W/m-K at 300K.B) Mass Transport Effects.andthe Growth RateWhile low-pressure LCVD processes are much too slow for macro-scale prototyping, the volumetric deposition rate rises with pressure. At pressures above 1 atmo~here, Wallenberger et aI. recently reported volumetric growth rates of nearly 0.1 mm3/s. 2 This rate increase occurs from greater adsorption of the precursor onto the reaction surface, as well as from enhanced natural convection--which permits more rapid transport to (and from) the surface. The growth rate continues to rise with pressure until the fluid reaches a critical pressure;26 for the metal halides, this critical pressure typically lies in the range of 25-100 atmospheres.

The growth rate continues to rise with pressure until the fluid reaches a critical pressure;26 for the metal halides, this critical pressure typically lies in the range of 25-100 atmospheres. Thus, by directing a forced flow of precursor fluids onto a heated reaction zone, an enhanced deposi tion rate is obtained, allowing parts to be grown quickly. This is the principle behind the HPCE-LCVD process. An added benefit is that the use of high pressures favors the growth of amorphous or fine-grainedpolycrystalline materials--which possess greater strength than cast, extruded, or sintered materials. These effects will be demonstrated later in this paper, where fine-grained, titanium rods were grown at enhanced rates.C) Chemical Vapor Deposition of TitaniumThe titanium halides are commonly employed for the chemical vapor deposition of tita nium.

These are pyrolyzed either directly, or by hydrogen reduction at temperatures above 1500 K.27 The overall reactions are:TiX4 (<J) --> Ti(s) + 2X2 (<J)'where X=CI, Br, or Ibrespectively. Titanium thin films have been deposited from TiCI4,28.29 as well as from TiBr4.3 Ti02 films have also been deposited31 from a mixture of TiCI4, H2, and 32 CO2--while titanium dioxide rods have been grown from TiCI4, H2, and atmospheric oxygen. II. ExperimentalA} Moderate Pressure Experiments (to 1.5 bar)The titanium growth experiments were carried out in a heated stainless steel reactor as shown in Fig. 1. The reactor (A) consisted of a six-way cross, with windows on two sides, one for laser input (B) and another for cross-illumination of the sample (not sh~wn). A UV-gra~e fused-silica window was located for observation at right angles to the laser Input and OpposIte the illumination source.

A UV-gra~e fused-silica window was located for observation at right angles to the laser Input and OpposIte the illumination source. A microscope and CCD camera were used to monitor the growth.Throughout the experiment, a 90 mm focal-length, gradient-index lens (C) was used to focus the gaussian input beam to a t/e2 spot size of approximately 50 microns. A Coherent Innova 70 argon-ion laser with output power to 2.2 W TEMoo (8 Watts Multimode) was used at498(1)Fig. 1: Apparatus for Moderate Pressure Experimentsthe primary 488/514 nm lines; this laser was readily interchanged with a Lee Laser 40 W (10 W TEMoo ) Nd:YAG laser. This was operated in cw mode at 1064 nm.A variety of substrates (D) were employed to initiate sample growth. Experiments were performed using bare titanium, glassy carbon, and graphite disks. The latter two were chosen for their low spectral reflectances and high melting points.

Experiments were performed using bare titanium, glassy carbon, and graphite disks. The latter two were chosen for their low spectral reflectances and high melting points. All substrates were mounted on a sample holder (E), attached to a micromanipulator, allowing the substrate to be scanned rela tive to the beam focus.Reactants were introduced into the reactor via a 1116" stainless steel entrance tube (F), which could be independently heated at the nozzle exit to temperatures up to 2400 K. Since titanium tetrachloride is a liquid at room temperature with a boiling point of 137 C, and tita nium tetrabromide and tetraiodide boil at 230 C and 377 C respectively, each precursor was first heated in a sample cylinder (G) to obtain a sufficient vapor pressure for growth. The sam ple cylinders were heated in a beaker of water or ethylene glycol (H).

The sam ple cylinders were heated in a beaker of water or ethylene glycol (H). During the static growth experiments, the precursor vapor pressure was prescribed by the temperature of this heated sample cylinder (G), while for the convective experiments, the roughing valve (I) was opened, allowing the overall chamber pressure to drop to less than 50 mbar; in this way, a constant flow was obtained down the 1116" stainless steel entrance tube (F), which could be preheated and directed toward the sample (D). The entire assembly was wrapped with heater tape and insula tion, except for the windows, to maintain a constant chamber temperature above that of the sample cylinder.

The entire assembly was wrapped with heater tape and insula tion, except for the windows, to maintain a constant chamber temperature above that of the sample cylinder. To eliminate condensation of the precursor on the laser window, a hot air gun (1) was directed at the window, at a temperature equal to or greater than that of the chamber.S) Design of the High-Pressure Experimental Apparatus (to 300 bar)In parallel with the apparatus described above, a high-pressure vessel has been constructed which will allow further investigation of metal deposition at much greater growth rates. The high-pressure chamber has been designed for operation up to 300 atmospheres, and has three fused-silica, Poulter-sealed windows, as shown in Fig. 2, allowing up to three orthogonal laser beams to be focused simultaneously to a common point (A). The high pressure reactor (B) is enclosed inside a heavy steel dome (C), which acts both as fumace and fume hood. Two499Fig.

The high pressure reactor (B) is enclosed inside a heavy steel dome (C), which acts both as fumace and fume hood. Two499Fig. 2: High Pressure Assembly and Reactor3000 W heaters circulate hot air around the reactor, allowing the windows to be warmed to the same temperature as the chamber. Laminated graphite and stainless steel gaskets are being employed to seal the windows and reactor cap, allowing chamber temperatures of up to 500 C. Precursor gases are supplied from a high-pressure evaporator (not shown), and are directed toward the laser focus (A) through a small nozzle. The chamber is evacuated using a simple Venturi pump, and reaction by-products are passed through a burn box to the clean room scrub ber.To scan the laser focus inside the chamber, four high-pressure motion feedthroughs are being installed in the lower portion of the reactor; these are attached to internal stages which position the sample and allow for three linear and one rotational degree offreedom.

In addition, two orthogonal beams will be modulated at the common focus, providing one more degree of freedom (tilt). In this way, completely arbitrary structures may be fabricated.III. Resultsthe room-temperature vapor pressure ofTiCl4 (Le. at abt. 13 mbar), no growth was observed on the glassy carbon substrates up to their softening temperature (abt. 1500 K). At higher pressures, the transient growth rate of each precursor improved steadily on the graphite and titanium sub strates. Samples (A), (B), and (C) in Fig. 3, were successfully grown at TiCI4, TiBr4 and TiI4 vapor pressures of 330, 8.5, and 180 mbar, respectively, and at transient rates of up to 0.37 p..mls. Titanium tetrachloride, with its greater vapor pressure, exhibited the highest growth rates, but only at temperatures approaching the melting point ofTi (1953 K).

Titanium tetrachloride, with its greater vapor pressure, exhibited the highest growth rates, but only at temperatures approaching the melting point ofTi (1953 K). To initiate growth with this precursor, we first ablated a portion of the titanium substrate, then lowered the aver age power (or defocused the beam) until the substrate could no longer be melted (10 W, 60 p..m spot size). Growth could then be initiated at incident powers as low as 4.3 W. Using the well-known expression for the peak temperature rise of a Gaussian beam on a semi-infinite solid (1):Experiments with the titanium halides atlow pressures were unimpressive. At500Fig. 3: Titanium Shapes from the Ti HalidesP(I-R) T = ,PJ'iKOlothis initial power indicates a TiCl4 threshold deposition temperature of approximately 1130 K.

At500Fig. 3: Titanium Shapes from the Ti HalidesP(I-R) T = ,PJ'iKOlothis initial power indicates a TiCl4 threshold deposition temperature of approximately 1130 K. Below this temperature no thick Ti films or rods were formed.At similar incident powers, titanium iodide yielded growth rates nearly commensurate with that of TiCI4, although at a fraction of the vapor pressure. It also exhibited a lower threshold deposition temperature of abt. 870 K (3 W incident power), which is consistent with reports that it possesses an activation energy lower than that of TiC14.33 The growth rate from TiBr4 was less impressive, giving rates of at most 0.04 p:m/s, at pressures up to 8.5 mbar.

Thermody namic calculations, minimizing the Gibb's free energy, confirmed that little degosition is likely to occur from pure TiBr4 at temperatures up to the melting point of titanium.3 However, fur ther experimentation at higher vapor pressures will be necessary to determine the potential of this precursor for HPCE-LCVD.Overall, titanium growth tended to be thermally self-limiting. This is illustrated in Fig. 4, where rods were grown for 30 minutes each from TiI4 at a variety of vapor pressures. Note that"8 200 -D ...., .j ! Po 100 dI l:iH160.jot'S.. J 60 j...080BO100 120 140 180 lB 0 200 220 240 280 2B 0 300 320 340'l'emperatule ( C)Fig. 4: Low-Pressure Transient Titanium Growth vs. Evaporator Temp.501(1)after initial rapid transient growth, each rod approaches a limiting height.

4: Low-Pressure Transient Titanium Growth vs. Evaporator Temp.501(1)after initial rapid transient growth, each rod approaches a limiting height. This can be explained either by rapid cooling as the deposit grows (and losses increase to its surroundings), or by overheating and melting as the rod grows away from the substrate (which acts diminish,.. ingly as a heat sink). To explore this self-limiting effect, the laser power was ramped up slowly, from an initial power of 6 W to over 10 W; surprisingly, this did not yield tall deposits as well as the opposite approach, Le. lowering the power from an initially high temperature. Thus, since insufficient convection occurs, the temperature increases6 as the rod lengthens, and the rod subsequently overheats and melts. Evidence of melting can also be seen in Fig. 3 (A), where the self-limiting rod height was attained.

Evidence of melting can also be seen in Fig. 3 (A), where the self-limiting rod height was attained. Clearly, steady-state6 rod growth cannot be obtained at low gas pressures unless the temperature of the deposit is lowered continuously in a controlled manner--and with the titanium halides little latitude exists between the tempera ture at which growth initiates and the melting point of Ti--making the reaction especially diffi cult to control.At higher vapor pressures, however, gas convection begins to contribute significantly to the heat losses, giving rise to a steady-state temperature distribution at the rod tip--and allow ing continuous 3-dimensional growth without temperature control. For titanium tetrachloride, this was realized at pressures above 2000mbar, as can be seeninFig. 5. At 3000mbar, Ti rods over 1 mm long were grown without changing the laser focus or the average power.

5. At 3000mbar, Ti rods over 1 mm long were grown without changing the laser focus or the average power. By gradu ally increasing the laser power as the rod grew (to compensate for the stationary laser focus and diminishing beam intensity), rods up to 4 rom long could be grown at axial rates up-to 1.4 pm/s. In addition, rods grown at the highest pressures exhibited no apparent grain structure, down to 100nm or less. This confirms the results ofWallenberger,25 who also obtained fine-grainedFig. 5: Titanium Rods Grown at High Pressure5024000BH .~ ....... m "8 2000 tri "8 .a .9 !30001000o OIeall Ex:Po&Ule Time (mi:Q1020304060o 10 OIeall i'.xPo&UIe Time (mi:Q203040Fig. 6: Rod Height and Diameter ys. Laser Powerand amorphous rods at high vapor pressures. In Fig. 6, the height and diameter of the Ti rods are also plotted vs. exposure time.

6: Rod Height and Diameter ys. Laser Powerand amorphous rods at high vapor pressures. In Fig. 6, the height and diameter of the Ti rods are also plotted vs. exposure time. As expected for steady-state growth, the rod height increases without bound, and the rod diameter approaches a constant (steady-state) radius.During the high-pressure reaction, a broad band of blue-violet emission lines were also observed, well below the blackbody emissions--which appeared to peak in the yellow portion of the spectrum. These have been identified as the strong 455-468 nm emission lines of the sin gly ionized Cl2+ molecule, a by-product of the TiCl4 reaction.Analysis of the Ti rod composition was carried out with an X-ray diffractometer. Combined results from several samples indicated that the rods were primarily pure Ti, with small quanti ties of Ti02• No TiCl3 was found to be present in the rods, although this solid by-product formed readily in a ring around the Ti rods.

The titanium rods were exposed to atmospheric oxygen prior to the X-ray analysis. Further measurements will be conducted using a scanning Auger microprobe to determine the extent of oxide contamination within the rods.IV. ConclusionsTo obtain continuous Ti growth, sufficient precursor pressure must be present for natural convection to establish a constant temperature distribution over the reaction zone. Without this convective heat loss, it is difficult to sustain growth; the temperature must be monitored contin uously to avoid melting the deposit. In addition, gas convection enhances transport of the pre cursor to the reaction, raising the volumetric growth rate by more than an order of magnitude.While steady-state growth may be obtained using TiCl4 at high pressures, the greatest promise for HPCE-LCVD may be TiI4, which exhibits the largest latitude between its threshold temperature and the melting point of titanium, facilitating growth control.

gh Quality, Fully Dense Ceramic Components Manufactured Using Fused Deposition of Ceramics (FDC)Cheng Dai, Gang Qi+, Sriram Rangarajan+, Suxing Wu+Noshir A. Langrana, Ahmad Safari+, Stephen C. Danforth+ Department of Mechanical and Aerospace EngineeringCenter for Ceramic Research-Department of Ceramic and Materials Engineering+ Rutgers - The State University ·ofNew JerseyAbstract Solid Freeform Fabrication (SFF) is a technology that produces physical solid components or parts from computer design models. This technology has the potential of reducing functional ceramic product development cycle time in terms of reducing design iteration and production time, minimizing extra post processing, and therefore reducing cost. A commercially available Fused Deposition Modeling (FDM™) 3D Modeler was altered for use with ceramics.

A commercially available Fused Deposition Modeling (FDM™) 3D Modeler was altered for use with ceramics. This newly developed method referred to as Fused Deposition of Ceramics (FDC) is capable of fabricating complex shape, functional ceramic components.We have investigated issues related to hardware, software, feed material, and build strategy which are required to achieve high quality, fully dense green ceramic parts. In this paper, we report recent improvements made in the FDC process, including hardware modifications, software improvements, feed material standardization, as well as build strategy/condition control. We also report the current FDC status for making complex functional parts. Our goal is to optimize the FDC condition to ensure its robustness for producing defect free green ceramic parts consistently and without interruption.1.

Our goal is to optimize the FDC condition to ensure its robustness for producing defect free green ceramic parts consistently and without interruption.1. Introduction: Solid Freeform Fabrication (SFF) is a technology that produces real physical solid models from 3D computer-aided design (CAD) model data, electronic scan data and model data created from 3D digitizing systems for rapid prototyping applications such as testing of form/fit/function. Because SFF is able to quickly go from 3D electronic data to an arbitrarily complex shaped 3D part, it allows design engineers to make design corrections very early in the product development process, and reduces the number of design cycle iterations and cost dramatically. The solid freeform fabrication technologies also have an advantage for small quantity production.

The solid freeform fabrication technologies also have an advantage for small quantity production. This is especially true for structural ceramic components, because the cost of manufacture of a structural ceramic component is a direct function of part complexity and production quantity.Fused Deposition of Ceramics (FDC), developed at Rutgers University, is an SFF technique [1 6], based on the commercial Fused Deposition Modeling (FDM) process for fabrication of ceramic and metal (FDMet) objects. FDC uses a polYmeric binder system to bond ceramic powders together in forming wire-like filaments which are melted in the liquefier delivery head which moves in the X and Y directions and is controlled by computer. The material is then411extruded from the heated head and deposited on a layer-by-layer basis.

The material is then411extruded from the heated head and deposited on a layer-by-layer basis. The extrusion process shears the material which quickly solidifies while reheating and bonding to the previous layers.For plastic and wax parts manufactured by the FDM process, the primary concern is defects on the part's surface. Defects such as internal voids have not been a top priority for FDM, because plastic and wax parts do not necessarily require full density for their applications. For ceramic parts made by the FDC process, internal voids due to build defects are crucial because they will act as critical flows that lower the fracture strength (i.e. functionality) of the ceramic component.Since our objective is to directly produce high quality functional ceramic parts, achieving full density in the FDC parts is the top priority in the building process.

The production of fully dense ceramic parts using FDC requires a high quality and continuous feedstock filament of FDC'able material and a fully automated FDC process. Functional ceramic parts also require good FDC process control of deposition tool path, build parameters and build environment.2. Build Strategy and Condition: Green ceramic parts can be built by FDC in different orientations. The decision of which build orientation to use depends on the part's geometry and application. Parts can be made horizontally or vertically, with support or without support. Parts also can be made with different selections of perimeter/contour/raster combinations, and with different angles of raster fill. Build conditions, such as liquefier and envelope temperatures are also important and need to be considered before part production.Producing high strength ceramic parts is a high priority in our FDC research.

The strength of a green FDC part will be affected by the build orientation and the build conditions used. Control of the part's environment during the build process is also critical for ensuring a high quality part. If temperature is not the build envelope temperature is too low or constant environment maintained, previously deposited material will cool too much and imperfect adhesion to the next layer will result. Incomplete inter-road bonding also occurs when tool path vector lengths are too long.Normally, a part with the largest cross section built in horizontal directions (X,Y) requires less build time compared with building the largest cross section in the vertical direction (Z). This is because of the Z-stage wait time of the 3D Modeler. When one layer is fInished, the Z platform has to move down to open the space for brushing the nozzle tip, then move upward to the build position. This action normally takes about 4 seconds.

When one layer is fInished, the Z platform has to move down to open the space for brushing the nozzle tip, then move upward to the build position. This action normally takes about 4 seconds. For example the tensile bar shown in Figure 1 required 0.6 hour when built horizontally and 1.3 hour when built vertically with the same head speed.Three RU955 GS-44 silicon niride tensile bars were made, all with 45, -45 degree raster fill shown in Figure 1. Two of them were made horizontally under different envelope temperatures, one with 45°C and one with 30°C.

Two of them were made horizontally under different envelope temperatures, one with 45°C and one with 30°C. One of the bars was made vertically with a 30°C envelope temperature.41245 degree raster fill45 degree raster fillFigure 1: Green GS-44 silicon nitride tensile test sample with 45, -45 degrees raster fills.According to the strength test results, Figure 2, the bar made in the horizontal direction with a 45°C envelope temperature had a higher strength (3.18 MPa) than the strength (2.84 MPa) of the identical bar made with a 30°C envelope temperature. However, the bar made in the vertical direction with a 30°C envelope temperature had the lowest strength (1.94 Mpa) of the three. The test data and stress vs. strain curves are shown in Table I and Figure 2. The results indicate that green FDC parts have a lower strength in the Z direction than in X-Y directions because of the nature of layer bonding in Z direction.

The envelope temperature has an effect on material bonding between layers, and residual stresses in the green part. The temperature be sufficiently high to ensure good bonding between layers. If the layers are cooled too rapidly the surrounding temperature), then residual stress may develop and lead to cracking during BBO (Binder Burn Out). In addition, if the temperature is too high, the part may deform under its own weight.should(duetoTable I: Tensile Test Results for Green RU955 Silicon Nitride Bars.EnvTemn 30°C 45°CRaster An!!le 45, -4545, -45Road Width 0.015" 0.015"STRENGTH Horizontal 2.84 MPa 3.18 MPaSTRENGTH Vertical 1.94 MPa4133. Sub-Perimeter Void and Software: The sub-perimeter void is a common defect for both FDM and FDC processes. The sub perimeter void occurs in the conjunction areas of perimeter and rasters. Because of the linear approximation of tool path, sharp turns for rasters occur.

The sub perimeter void occurs in the conjunction areas of perimeter and rasters. Because of the linear approximation of tool path, sharp turns for rasters occur. As a result, empty spaces (voids) are left without being filled due to the curvature of the perimeter and sharp comers (angles) of rasters. These sub-perimeter voids in FDC parts are too large to be overcome in the sintering process, hence the strength of FDC ceramic parts is affected significantly.Recently, we have been using a simple method to eliminate sub-perimeter voids in the FDC process. This method might not be the best solution for solving this problem, but it is an easy and effective way to get ride of the sub-perimeter voids.Normally, when creating the tool path on each slice, the build parameters need to be specified in the software QuickSlice. The air gap between roads can be set with negative, zero, or positive offsets depending on the application.

The air gap between roads can be set with negative, zero, or positive offsets depending on the application. For example, if the offset is set to a positive offset, then there is a real gap between fill vectors. If the air gap is equal to zero, then there is no air gap, which in principle means perfect filling and that the roads are perfectly in contact with each other. If the air gap has a negative offset, there will be overlapped fill. In the past, we have used small negative offsetting of the air gap to ensure internally fully dense parts. In fact, the function of the offsetting technique can be used effectively for the elimination of sub-perimeter voids.

In fact, the function of the offsetting technique can be used effectively for the elimination of sub-perimeter voids. This simple method works well in most cases especially for our RU955 silicon nitride parts.ABFigure 3 : X-ray radiographs of green GS-44 silicon nitride part made with a negative offsetting (A), and without negative offsetting (B).Figure 3A shows an X-ray radiograph of a FDC green GS-44 silicon nitride part sub containing perimeter voids. This was accomplished by specifying an air gap with a negative offsetting in the build The parameters. negative offsetting allowed to overlap the raster fill with fill, eliminating the possibility of voids. However, whennothecontouran air gap with no offsetting was specified, sub-perimeter voids were created.

However, whennothecontouran air gap with no offsetting was specified, sub-perimeter voids were created. Figure 3B shows an X-ray radiograph of a part, built without a negative offset, containing voids.Overflow (too much material deposited on the part surface) created by using the negative offsetting is the disadvantage of this method. It sometimes causes dimensional changes in parts.414However, by carefully selecting the value of negative offsetting, the overflow problem can be minimized. We have been using negative offsetting from -1/4 to -1/3 of the raster width in most cases. The dimensional expansion of the part sometimes can be as large as 0.008", depending on the build condition and part geometry.This negative offset technique works well for rasters with small road widths. This is because of the void created at the sharp turning comer for small road width rasters is smaller than the one created by large road width rasters.

This is because of the void created at the sharp turning comer for small road width rasters is smaller than the one created by large road width rasters. For parts which have an interior small boundary, e.g., a small hole, it is difficult to eliminate the sub-perimeter voids especially when the size of a small boundary or hole is not much greater than the road width of the raster fill.4. Feed Material Standardization: In the past, we had frequently experienced clogging of the nozzle in the FDC process especially for small size nozzles (0.020"D or less). The causes of nozzle clogging were foreign objects and agglomerates of ceramic particles. Recently, the FDC filament fabrication laboratory at Rutgers University has gone through a series of steps for improving the quality of the FDC filament [8, 9]. The main focus was to eliminate any foreign contamination, and to produce better mixing/deagglomeration and more uniform filament.

The main focus was to eliminate any foreign contamination, and to produce better mixing/deagglomeration and more uniform filament. A fine screen is used in filament extrusion to remove large particles or foreign objects. The Standard Operating Procedure (SOP) of filament fabrication has been implemented to ensure reproducibility and to minimize the chance of contamination and process variation.Filament buckling is another critical problem in the FDC process. Buckling is caused by poor quality of filament and insufficient filament aging. Poor quality filament refers to filaments which either are not stiff enough, or have high viscosity, or is not uniform in diameter. Most buckling due to poor filament quality has been eliminated via improved ceramic processing. In most cases, we believe that the filament buckling is due to a material aging problem. In the recent past, we experienced some difficulties due to the inconsistency of FDC feedstock material.

In most cases, we believe that the filament buckling is due to a material aging problem. In the recent past, we experienced some difficulties due to the inconsistency of FDC feedstock material. A large part of the problem was due to the effect of filament storage and due to the lack of understanding of the RU955 binder's aging. In order to produce the same condition of feed material which is feasible to FDC on a daily base, we performed a series of aging tests for the RU955 binder [8, 9].According to the test results, moisture does affect the quality of RU955 filament. Too high a moisture content of the FDC filament results in a high drive motor current and bubbles on the extruded surface. However, the aging studies suggest that the moisture effeyt can be eliminated by a vacuum heat treatment as part of filament aging. A fixed and standardized filament aging process and storage procedure has been employed to ensure consistent FDC conditions [8, 9].5.

A fixed and standardized filament aging process and storage procedure has been employed to ensure consistent FDC conditions [8, 9].5. Hardware Improvements: Since the filament used in the FDC process also acts as the piston for extrusion, the filament should have enough stiffness when entering the liquefier. It has been shown that the G' (storage modules of the filament) is a strong function of temperature between OOC and 30°C. Therefore, cooling the filament near the top of the liquefier becomes critical. If the heat brought up from the415liquefier is not removed, then the heat will soften (lower the G' - stiffness) the filament and in tum cause buckling. Because the cool air (at R.T.) from the FDM cooling system at the top of the head is partially blocked by the wheels and motor block, the FDM cooling system is insufficient to maintain enough stiffness in the FDC filament.

The cooling path shown in Figure 4 for FDC has been altered from the top to the right side to provide cooling directly in the region where the filament can buckle. In the new cooling system, the cooling path and the filament feeding path have been separated above the work platform. A temperature controlled cooler is attached to the air blower on the back of the machine, so that cool air (-15°C-18°C) is blown by the cooling fan into the liquefier head. The air then goes back to the cooler's thermostat chamber to make a closed-loop so that the temperature can be controlled.

The air then goes back to the cooler's thermostat chamber to make a closed-loop so that the temperature can be controlled. This modification has dramatically reduced the filament buckling problem, allowing better process control, and maintenance of more uniform FDC conditions.Temperatures under three different conditions were measured for one FDC build condition (liquefiertemperature=186°C; envelope temperature=44°C; envelope humidity=25% RH), they are with AlC (-18°C) cooling, without AlC but with room air (-25°C) cooling, and no cooling. Temperatures for each case were measured at two locations, at the location between filament rollers and at top of the liquefier (see Figure 4). The measured temperatures are listed in Table II,a e ~ hZ 11 ~emperatureeasurementsTemp. Between Rollers r'C)Temn.

The measured temperatures are listed in Table II,a e ~ hZ 11 ~emperatureeasurementsTemp. Between Rollers r'C)Temn. at Ton of Liauefier COC)Wi'th l No Coolingand Wi'thl out C Z'00 mg.25°C RT Coolin!! 28-29 32-3543-4561-6418°C AlC 20-21 24-25The thermostat on the AlC unit was set at 18°C, and it continuously pumped 18°C cool air into the liquefier head. The cool air then leaves the head from the electric wire conduit to the back of the 3D Modeler and back to the air conditioner. Since it is a closed-loop and the cooling system build does not envelope the envelope temperature can still be well controlled. The FDC build envelope temperature recently has been raised to 45°C from previous values used 32°C-38°C.

For example, the port plate shown in Figure 5(a) has several cavities. Improper selection of road width will cause internal fill air gaps most likely appearing in two regions. One is the area where the raster fill is parallel or near parallel to a curved boundary (sometimes, called high angles). The second is the area where two boundaries are separated from each other by a small distance. Improper selection of road width will cause voids in these two regions.As an example, because of the complex geometry of the port plate shown in Figure 5(a), it is necessary to use more than one "set" of build parameters for creating internal roads. The advantage of using multiple sets for creating internal roads is that each set will match well with assigned slices which represent a particular geometry. The multiple assignments of sets will overcome the difficulty for handling a complex part by a single set, and will eliminate internal iterations of internal road design. It defects.

The multiple assignments of sets will overcome the difficulty for handling a complex part by a single set, and will eliminate internal iterations of internal road design. It defects. However, complex objects normally require several is therefore important to investigate all different slices of internal roads created by QuickSlice before saving it into a build file or download it to the 3D Modeler.The port plate shown in Figure 5(a) designed by AlliedSignal has a complicated geometry with different size holes and under cuts which is very difficult and expensive to make by normal manufacturing processes. A total of 3 port plates have been made by FDC recently. The turbine blade for an auxiliary power unit (APU) shown in Figure 5(b) is also a complex part because of the twisting curvature of the blade. Recently, we made several 1.5" tall turbine blades using the FDC process.

Recently, we made several 1.5" tall turbine blades using the FDC process. The overall dimension, FDC build speed and time, and length of filament used for each port plate and turbine blade are listed in Table III.(a)(b)(c)(d)Figure 5: Demonstration ofrecent green FDC parts made out ofsilicon nitride.The human head model and the curved cylindrical shell shown in Figure 5(c) and 5(d) respectively, were made by using a regular FDM O.015"D nozzle. It took 4.6 hours build time to417make one human head model with 0.007" slice thickness at O.5"/sec speed. During the FDC process of these parts, there was no human intervention and no process interruption. It has been shown that the current FDC process has great potential and the flexibility to use small size nozzles (0.015"D) and therefore improve the part surface quality.

It has been shown that the current FDC process has great potential and the flexibility to use small size nozzles (0.015"D) and therefore improve the part surface quality. The surface quality of the parts made out of the 0.015"D nozzle are much better than using 0.025"D nozzle, and in some applications, may not require further machining.Table III: Data Sheet for Each ofthe Port Plates, Turbine Blades, and the Radome and Cellular Filter Made Out ofGS-44 Silicon Nitride.FDCPartPort Plate Turbine Blade Radome Cellular FilterFDCPart Dimension2.7"DxO.6"tall 1.5" tall 4.5"Dx7.5"tall 3"DxO.5"tallFDC Build Speed (in/sec)0.5 0.5 0.5 0.5RU955 Si3N4 Filament (rt) 65 24 117 41FDC Build Time (hour)9.5 4 16 4.3Among all of the parts made recently, the largest part is the section of a missile radome shown in Figure 6 designed by Lockheed Martin. The radome section is 7.5" tall and the base section diameter is 4.5".

The radome section is 7.5" tall and the base section diameter is 4.5". This radome was made in 16 hours at O.5"/sec speed and used approximately 117ft of FDC filament. It was manufactured with total computer control in a fully automated process. There was not a single interruption during the 16 hour run. The RU955 GS-44 Si3N4 filament was used just like the ICW-05 wax filament, i.e. it was on a spool and automatically and continuously fed from the back of the 3D Modeler. During the entire process the Modeler's door was kept closed and the build envelope temperature was controlled at 40°C. A O.015"D nozzle was used and there was no nozzle clogging. The surface quality of the radome is excellent even without green fInishing.

A O.015"D nozzle was used and there was no nozzle clogging. The surface quality of the radome is excellent even without green fInishing. So far, this is the longest uninterrupted FDC build since the system has been modified.Figure 6: Green silicon nitride Lockheed Martin's radome made by FDC.Figure 7: Green silicon nitride cellular filter substrate made by FDC.418Another complex part that has been made is the cellular filter substrate, which was designed and made for use as a sintering base as shown in Figure 7. This part is 3" in diameter and 0.5" thick. Because of its fine mesh structure, it is difficult and expensive to manufacture by the standard commercially available manufacturing processes. The filter was made in a 4.3 hour FDC build at 0.5"/sec speed with a 0.015"D nozzle, without interruption. The mesh structure was created by specifying a positive 0.040" offset between raster roads for the internal fill of the disk.

The mesh structure was created by specifying a positive 0.040" offset between raster roads for the internal fill of the disk. The green cell wall thickness was 0.025", and the overall part quality was excellent.7. Conclusion: The success of the FDC manufacturing process demonstrated in this paper is mainly due to the recent efforts on improved filament fabrication, feedstock material standardization, new FDC build strategies, hardware modifications, and improved process control. The degree of automation for high quality FDC process has been increased dramatically. Now, the FDC process is capable of making complex, fully dense, and high quality ceramic parts using a fully automated FDC process. Since January, 1997, we have made approximately 100 parts using FDC without human intervention or interruption.

Since January, 1997, we have made approximately 100 parts using FDC without human intervention or interruption. This success has demonstrated the capabilities of the improved FDC process.Further improvements in the areas of elimination of material flow variation, elimination of contamination, and good tolerance (shrinkage) control need to be made. Increasing the build speed along with using a fine sized liquefier nozzles need to be investigated further. Better process control and improved surface finishing are needed for commercialization [10].Acknowledgments: We gratefully acknowledge the support of this research by DARPA/ONR under contract #N00014-94-0115. This research has been conducted in the Laboratory for Solid Freeform Fabrication of Advanced Ceramics in the Center for Ceramic Research and the Department of Mechanical and Aerospace Engineering at Rutgers University, and the Research and Technology, AlliedSignal.The authors wish to thank Dr. P. Whalen, Dr. V.

P. Whalen, Dr. V. Jamalabad, and Mr. R. Clancy of AlliedSignal for their valuable suggestions and contributions to this research. The authors also would like to thank Dr. P. Bhargava, Dr. A. Bandyopadhyay and Mr. R. C. McCuiston of Rutgers University, and Dr. G. Carrasquillo of Certech, Inc. for their help and contributions to this work.Reference: 1.A. Bandyopadhyay, S. C. Danforth, A. Safari, C. Dai, G. Qi, S. Rangarajan and N. A. Langrana, "Structural Silicon Nitride Components via FDC", The American Ceramic Society's 99th Annual Meeting and Exposition, May 4-7, 1997. M. K. Agarwala, R. van Weeren, A. Bandyopadhyay, A, Safari, S. C. Danforth, W. R. Priedeman, "Filament Feed Materials for Fused Deposition Processing of Ceramics," Proceeding of the Solid Freeform Fabrication Symposium 1996, pp. 451-458, Austin, Texas, August 12-14, 1996, University of Texas at Austin.2.41934.5.6.7.8.9.10.V. R. Jamalabad, N. A. Langrana, and S. C.

gh Resolution Laser Melting with Brilliant RadiationA. Streek, P. Regenfuss, H. ExnerLaserinstitut der Hochschule Mittweida, Mittweida, D-09648KeywordsLaser micro sintering, metals, additive manufacturing, powder, densityAbstractSince the discovery of selective laser sintering/melting, numerous modifications have been made to upgrade or customize this technology for industrial purposes. Laser micro sintering (LMS) is one of those modifications: Powders with particles in the range of a few micrometers are used to obtain products with highly resolved structures. Pulses of a q-switched laser had been considered necessary in order to generate sinter layers from these µm-scaled metal powders. However, despite the high resolution, the process repeatability of LMS and the material property of the products have never been completely satisfactory.

However, despite the high resolution, the process repeatability of LMS and the material property of the products have never been completely satisfactory. Recent technological and theoretical progress and the application of brilliant continuous laser radiation have now allowed for efficient laser melting of µm-scaled metal powders. Thereby, it is remarkable that thin sinter layers are generated with a very high laser power. The resulting product resolution is comparable to the one achieved by the LMS regime with q-switched pulses. From the experimental results the performance and potential of this high resolution laser melting regime is demonstrated and the limits of the applicable parameters are deduced.IntroductionLaser micro sintering is a modification of selective laser sintering, which is based on repeated cycles of coating and selective densification (sintering) of powders or of pasty materials, with the ends of generating, layer by layer, a three-dimensional body.

In accordance with the name, the solidification of a product layer is achieved by scanning a laser beam over the powder surface within the cross-section of the desired object. The material passes through a transitory melt phase. Since its invention by Carl Deckard & colleagues [1] selective laser sintering has functional been upgraded continuously to meet components [2]. A considerable contribution to the improvement of resolution was the modification and optimization of the already established technology in 2001 by Laserinstitut der Hochschule Mittweida. It resulted in a technology termed laser micro sintering. The main features are the employment of q-switched laser pulses and the use of powder with particle sizes in the µm-range. Resolutions have been achieved that approximate the focal diameter (25 µm). A great disadvantage was the limited powder bed density of the applied fine grained powders.

Resolutions have been achieved that approximate the focal diameter (25 µm). A great disadvantage was the limited powder bed density of the applied fine grained powders. Therefore, the recoil forces of the expanding plasma generated by the high intensive q-switched pulses were necessary as a powder condensing momentum. Because of limited intensities of commercially available laser sources in the past a technique for laser sintering of fine grained powders with a cw-Laser had never been accomplished. Applying new high power laser sources with highest brilliance coupled with a powder compacting mechanism allows for laser micro sintering with cw-radiation, which is commonly designated “laser melting”.the requirements forthe production ofExperimental setupThe overall experimental setup for laser melting is principally comparable to the one for laser micro sintering except for a new condensing powder coater [3] and a polygon scanning device.

The powder coater was developed to produce highly accurate and dense powder beds of the employed micro powders (Fig. 1.).Fig. 1: (a) The novel powder coating forces the powder under high pressure, to flow from the reservoir onto the building platform, to produce highly condensed and homogeneous powder beds of the respective thicknesses. (b) Close view showing the coating properties of the applied molybdenum micro powder.A YLR-400-LP-AC single-mode fiber laser with a nominal laser power of 400 W from by IPG Laser GmbH (Burbach, Germany) was applied in the continuous wave mode. A 230 mm f-theta- optic, attached to the scanning device, focuses the beam with a gaussian intensity distribution and a radius w86 of 35 µm in the plane of the powder bed surface. For scan velocities up to 10 ms-1 a SuperScan-LD-30 from Raylase AG (Weßling, Germany) is used.

3).Fig. 3: The applied molybdenum powder Mo3-7. (a) SEM-view of the powder material. As denoted, the grain distributions by mass are d10 = 3 µm and d 90 = 7 µm. Therefore the largest number of grains belongs to the fraction of smaller particles and is partially agglomerated. (b) The nominal powder density and therefore the coating density obtained with a classical additive manufacturing setup is 20 %. With the novel powder coating device a compaction of the powder above a density of 30 % is possible.BackgroundIn a stable laser melting process, the applied process parameters have to allow for a minimum melt depth in the order of the thickness of the coated powder layer and furthermore for effective wetting of the underlying sinter layer. Therefore, the absorbed energy per volume has to suffice for the fusion of the material, which includes the heating to melting point TM (and above) and the latent heat for the phase transition.

Therefore, the absorbed energy per volume has to suffice for the fusion of the material, which includes the heating to melting point TM (and above) and the latent heat for the phase transition. However, the absorbed energy has to be lower than the amount, at which the overall ablation, including the expulsion of already molten material, exceeds the material mass added to the site by each powder coating cycle. As already known from investigations of laser micro sintering by high intensive q-switched pulses [5], this will not necessarily lead to a breakdown of the building process but certainly to a noticeable increase of the “height defect” and to deterioration of the micro part resolution. The resulting melt depth ℓm of a massive material (infinitive rod) can be transformed to the simple Eq.

The resulting melt depth ℓm of a massive material (infinitive rod) can be transformed to the simple Eq. 1 by the solution of Fourier’s equation for one-dimensional heat conduction, under consideration of the thermal diffusion length ℓD, the room temperature TR and an exponential increasing thermal heat gradient within the material, resulting from a steady increase in surface temperature TS by the absorbed laser radiation as suggested in [6] and further analyzed in [5].mln   T  TSM T R TR   DTSTMThe maximum surface temperature TS is reached when it equals the boiling point TB. This limitation defines the highest intrinsic energy that the material can bear without boiling and thus the maximum irradiation time as well as the corresponding intensity. For various materials the dependence of the maximum melt depth on the irradiation time is shown in Fig. 4.(1)Fig.

For various materials the dependence of the maximum melt depth on the irradiation time is shown in Fig. 4.(1)Fig. 4: Achievable melt depths ℓm, of different massive metals at irradiation times up to 1 ms under the condition that the material specific boiling temperature is reached at the end of the irradiation. The column on the right lists the elements in the order of their melt depths.For the investigated molybdenum, under ideal conditions an irradiation time of 50 ns is necessary to achieve a 1 µm depth of the melt pool. The (logarithmically) scaled diagram shows a steady increase of the achievable melt depth along with the irradiation time. Following these simplified estimations, after 500 µs, which means a maximum scan velocity of 0.1 ms -1 for a focused 50 µm laser beam, 100 µm melt depth could be achieved. By applying larger focus diameters increased scan velocities can be chosen to achieve the necessary melt depth, but logically the resolution will deteriorate.

By applying larger focus diameters increased scan velocities can be chosen to achieve the necessary melt depth, but logically the resolution will deteriorate. As already originally discussed, in [7], overall energy absorption within and energy distribution throughout the powder bed is more complex than the above formulated approximation. Another shortcoming, especially with small powders, is that most often the thermal diffusions length surpasses the particle diameter. In that case, the calculation based on the assumption of an infinite rod will yield an incorrect melt depth. This fact is principally demonstrated by a numerical temperature field calculation for three laser irradiated powder particles with different radius r [5] with consideration of material expulsion by vaporization and of the latent heats. I0 = 2.17 Wcm-2 (Fig. 5: data) is the minimum required intensity to achieve the boiling temperature at the surface of an infinite rod after 180 ns.Fig.

Boiling and loss of material occurs in the smallest particle (1:r = 0.6 µm) due to virtually-zero heat dissipation. Later on, in Fig. 8, these calculated cases 1-3 will be represented with likewise numeration as green circular dots.The increase of the melt depth in particles with dimensions smaller than the thermal diffusion length is obvious. Therefore, smaller particles should allow for laser melting with higher scan velocities at smaller focus diameters. Furthermore, due to lack of heat dissipation into deeper layers, thorough melting of the particle can be achieved after the laser irradiation, if the overall transferred energy during the irradiation time is larger than the thermal energy, required to raise the temperature to the boiling point plus the necessary amount of latent heat for the melting of the entire particle mass (or of the residual mass, in case of loss through boiling).

For a finite rod this critical depth dℓm for thorough melting is analytically described in [5] resulting in Eq. 2,dm   TMTRH cpmTB  1TR dmm        1e2dmD    2 th cpHwith ΔHm the specific latent heat of fusion, cP the specific heat capacity, λth the thermal conductivity, and ρ the specific density. This critical depth is the maximum depth of thorough melting if the boiling temperature is reached at the end of the irradiation time. For any other final surface temperature TB has to be substituted by TS (maximum surface temperature). In Fig. 6 the maximum critical depths are presented for different materials and irradiation times.(2)]m µ [403020Tungsten Iron TitaniumMolybdenum Copper Aluminiummℓd10000,511,5τH [µs]Fig.

6: Critical depths for thorough melting of a finite rod after a certain irradiation time. Due to the large heat conductivity of aluminum and copper these materials also show the deepest thorough melting after short-time laser irradiation. In tungsten and molybdenum a larger amount of energy is storable. Therefore these elements range above iron and titanium with the smallest heat conductivity. 300 K was chosen as environmental temperature TR in the calculation.For molybdenum this means that with 1 µs irradiation (velocity ~ 50 ms -1 @ about 50 µm spot size) the maximum achievable melt depth beneath the surface is almost doubled from 5 µm in bulk material to 10 µm in an isolated particle.

As already mentioned, to achieve the maximum melt depths a specific amount of energy, respectively a specific intensity of the laser, is necessary to just reach the boiling temperature at the end of a chosen irradiation time, at which, with this optimum intensity Iopt, the maximum energy storability is reached. Under the same assumptions as for Eq. 1, for a bulk material (infinite rod) the optimal intensity can be analytically described by Eq. 3,Iopt th D  TBTRlnT TSM T TRRH cPm  Iopt1 Hwhich clearly shows the inversely proportional relation between the applicable intensity and the irradiation time τH (included in ℓD).

Additionally, an intensity window can be defined for laser melting, between a minimum intensity at which the surface temperature just reaches the melt temperature, and a maximum intensity at which an additional amount of latent heat for boiling to the depth of ℓm is inserted and most of the material vaporizes. This intensity window (between the dashed lines) and the optimum intensity for different materials are presented in Fig. 7.Fig 7: Dependence of the optimal intensity (solid line) and the boundary intensities for different metals on the irradiation time up to 100 µs. The plotted intensities represent the absorbed intensities; the specific absorbance has to be considered for each material.2(3)For the definition of a process window for particles that are considered finite rods, two dependencies have to be observed: Firstly the intensities that lead to a certain temperature after thermal equilibration of the particles with various diameters. These dependencies are linear.

These dependencies are linear. Secondly the intensities that effect a certain temperature on the surface of the particles with the end of the irradiation time. These dependencies have a curved various diameters at three equilibrium progression. temperatures and for three surface temperatures respectively. The straight lines in Fig. 8 consist of the intensities by which, disregarding possible over-heating, the cross section of the respective particle absorbs, during the pulse of 180 ns, respective characteristic amounts “min(Q)”, “opt(Q)”, and “max(Q)”. “min(Q)” of energy. “min(Q)” would be just sufficient to convert the entire particle into liquid phase. “opt(Q)” displays the intensities, at which the enthalpy of fusion and additionally the energy that is required to raise the particle temperature to the boiling point is absorbed. This is the intensity at which the entire particle can be converted into a melt phase with boiling temperature.

This is the intensity at which the entire particle can be converted into a melt phase with boiling temperature. “max(Q)” is the respective minimum energy that is needed to raise the temperature of the entire particle mass to the boiling point, plus the energy for comprehensive evaporation of the particle.In Fig. 8 both types of dependencies are displayed forFig. 8: Characteristic (absorbed) intensities for the fusion of molybdenum particles with a 180 ns laser pulse vs. particle radius. The straight lines are the intensities that yield per pulse the respective net amounts of energies per particle cross section, that would be sufficient to fuse the particle (min(Q)), to fuse and heat the particle to boiling temperature (opt(Q)), and to heat the particle to boiling temperature and evaporate the entire mass (max(Q)) in an imaginary, slow heating process.

The bent curves consist of the intensities absorbed from a 180 ns laser pulse that lead to the fusion temperature on the irradiated particle surface (min()), the boiling temperature on the irradiated particle surface (opt()), and to comprehensive boiling of the fused portion of the particle when the irradiated surface has reached the boiling temperature (max()). The green dots represent the chosen parameters of the thermal simulation in Fig. 5.The second prerequisite is determined by heat conduction of a finite rod described in [8]. As a first approximation the heat conduction in spheric particles has been equated with the one in cylindrical rods of the same cross-section and volume; on the abscissa, the radii of the spheric particles are indicated. The three intensity curves in Fig.

The three intensity curves in Fig. 8, max(),opt(), and min(), represent the following thermal conditions at the end of the pulse: Curve “min()” consists of the intensities at which the particle surfaces have reached the melt temperature at the end of the pulse. Any further energy would lead to the onset of melting in the surface zone. It is considered the minimum energy where partial fusion of particles might be observed. “opt()” are the intensities at which the irradiated surface acquires boiling temperature. With these intensities, the maximum energy can be absorbed by the respective particles without any ablative loss.With intensities of curve “max()” not only the irradiated surface reaches boiling temperature, but additional energy has also been absorbed that is equivalent to the evaporation enthalpy for the already fused portion of the particle mass. This would imply ablation of the entire liquid phase and preclude any inter-particular fusion.

This would imply ablation of the entire liquid phase and preclude any inter-particular fusion. The shaded section between the curves “min(Q)” and “opt()” denotes the range of particles that can be comprehensively converted into liquid phase without ablation by a 180 ns pulse with the respective intensity. The largest molybdenum particle (max(rp-melt) in Fig. 8) that can be comprehensively fused has a radius of around 2.1 µm. Larger particles require longer irradiation times (Fig. 9).Fig. 9: Intensity/particle-size windows for comprehensive melting of molybdenum particles with laser pulses of various pulse widths, deducted analogue to the values in Fig. 8.ExperimentalFor laser micro sintering with q-switched pulses the maximum achievable scan velocity was defined by the given pulse repetition of the laser.

8.ExperimentalFor laser micro sintering with q-switched pulses the maximum achievable scan velocity was defined by the given pulse repetition of the laser. For highly dense micro parts the maximum pulse distance (at a beam diameter of 30 µm) was about 10 µm, resulting in scan velocities below 1 ms-1 for the applied 20 W q-switched laser with a repetition rate of 80 kHz. The necessary absorbed average intensity for the applied molybdenum powder was about 2∙107 Wcm-2 for 180 ns irradiation time, assuming a material absorbance of Amat = 0.30 from the fresnell equations (Fig. 10).Fig. 10: Resulting sinter densities from molybdenum at various pulse distances and intensities [Wcm-2].

10).Fig. 10: Resulting sinter densities from molybdenum at various pulse distances and intensities [Wcm-2]. The highest densities are achievable at 5 – 6 Wcm-2 (irradiating intensities) and at a pulse overlap of 2-3.Under consideration of the necessary pulse overlap, the energies for high density micro parts are 2-3 times higher than theoretically estimated (at 2 Wcm-2 only one single pulse should suffice to thoroughly melt the particle). Significant increase of the density at higher pulse overlap (at pulse distances around 10 µm) could only be observed if the pulses were set in a continuous sequentialalignment (Fig. 11). Otherwise the densities were almost equal to those achieved with high pulse distances. The remanent heat from the previous pulse seems to support the wetting of the melt. With a cw laser this behavior could also be expected at lower scan velocities that allow for sufficient accumulation of heat.Fig.

With a cw laser this behavior could also be expected at lower scan velocities that allow for sufficient accumulation of heat.Fig. 11: Sinter structures resulting from (a) stochastic and (b) sequentially line wise pulse positioning. The density of the applied pulses is the same for both specimens. The resulting sinter density was (a) 66 % and (b) 97 %. Sinter parameters: IAv = 6.5∙107 Wcm-2, w0 = 12,5 µm, τH = 180 ns, fP = 80 kHz, Molybdenum.Based on the theory, irradiation times of 2.4 µs for isolated particles and around 10 µs for bulk material should be sufficient/necessary to melt a 10 µm sinter layer, if an absorbed intensity between 0.3∙106 Wcm-2 and 0.62∙106 Wcm-2 can be reached. With the 400 W cw laser a maximum intensity of 4∙107 Wcm-2, respectively an absorbed intensity of 1.2∙107 Wcm-2 by the molybdenum powder is achievable at 35 µm spot size.

With the 400 W cw laser a maximum intensity of 4∙107 Wcm-2, respectively an absorbed intensity of 1.2∙107 Wcm-2 by the molybdenum powder is achievable at 35 µm spot size. With the attached galvanometer scanner a maximum scan velocity of 8 ms-1, respectively a minimum irradiation time of 4.4 µs could be achieved resulting in the sinter densities presented in Fig. 12.Fig. 12: Resulting sinter densities at various intensities for seven irradiation times corresponding to scan velocities from 2 – 8 ms-1. For comparison, the maximum achievable melt depths ℓm calculated with equation 1 are listed.With a (too) long irradiation time of 17.5 µs high sinter densities can be achieved throughout the entire time range. At short irradiation times (respectively with melt depths that will not considerably reach beyond an isolated particle) the resulting density decreases significantly. However, the determined sinter densities in Fig. 12 do not reflect the achieved accuracy of thespecimens.

However, the determined sinter densities in Fig. 12 do not reflect the achieved accuracy of thespecimens. At high intensities and long irradiation times the accuracy of the sintered micro parts deteriorates significantly (Fig. 13: top). The specimens obtained with low intensities and short irradiation times had a porous consistency. Loss of accuracy may be due to poorer fusion of the material in the peripheral zones (Fig. 13: bottom).Fig. 13: Cross section preparations of specimens generated at various intensities and irradiation times by a cw laser. Top: 2ms-1 (17.5 µs). Bottom: 8 ms -1 (4.4 µs).Furthermore, the specimens generated with short irradiation times and low intensities (Fig. 13) showed a behavior of sintering that is familiar from the stochastic pulse distribution (Fig. 11(a)). As expected, specimens with these parameter sets are generated with a noticeably insufficient wetting of the underlying (previously generated) sinter layer.

11(a)). As expected, specimens with these parameter sets are generated with a noticeably insufficient wetting of the underlying (previously generated) sinter layer. At ideal irradiation times and intensities well defined micro structures at high densities could be generated (Fig. 14).Fig. 14: Surface structures of a laser melted grid. Compared to laser micro sintering relatively thick sinter layers of 10 µm with could be generated with high resolution. This accelerates the building rate significantly compared to the laser micro sintering with merely 1 µm sinter layer.The resulting roughness of the micro structures was about 35 µm with a final accuracy of about 60 µm for free standing walls (Fig. 15).Fig. 15: (a) Surface morphology of the laser melted specimens. (b) Free standing walls with different angles.With higher scan velocities above 15 ms-1 that can be achieved with the polygon scanner device no laser melting of 10 µm thick sinter layers was possible at all.

Although theoretically, the critical melt depth for isolated particles at the corresponding irradiation time of about 1 µs should be sufficient (Fig. 6), the melt will never wet the underlying sinter layer. Upon considerable reduction of the sinter layer thickness down to 1 µm, which is less than the achievable melt depth for the bulk material (Fig. 4), the first solid results could be observed (Fig. 16). This proves the feasibility of laser melting with high scan velocities.Fig. 16: Molten tracks of molybdenum powder at various high scan velocities and the respective short irradiation times. The applied cw-laser intensity was limited by the maximum power of the 400W laser to 4∙107 Wcm-2.The decline of the fusion with the underlying sinter layer is apparent. At the highest scan velocities the resulting melt fusion occurs mainly in between the particles (Fig.16(c)), which is in the order of the calculated melt depth of about 1 µm.

At the highest scan velocities the resulting melt fusion occurs mainly in between the particles (Fig.16(c)), which is in the order of the calculated melt depth of about 1 µm. For noticeably better vertical fusion the irradiation times have to be 2-3 times longer than estimated for the respective sinter layer thickness. This corrected irradiation value approximates quite well the time that is required to obtain a critical melt depth in bulk material that equals the sinter layer thickness. Furthermore, this relation corresponds to the necessary 2-3 - fold pulse overlap, and thus to the 2-3 - fold energy excess, that is required for the achievement of higher densities in laser micro sintering (Fig.

10).SummaryIn the last years, as a consequence of improvements in the powder coating technique, it became evident that highly resolved laser sintering with small powders does not necessarily imply abandonment of the high sinter densities that are achieved in laser melting with larger particles. With the available higher powder densities it had become possible to generate solid bodies with a line-wise strategy whereby remanent heat can be exploited. It was already anticipated that with tightly focused lasers that would supply sufficiently intense cw-radiation, rapid laser melting with high resolution would be feasible without the employment of q-switched pulses.

Systematic experimental investigation has now delivered the proof that selective laser melting with resolutions comparable to the ones achieved in laser micro sintering can be executed with brilliant cw-laser-radiation.The recently developed model of a crucial range concerning the combination of intensity, exposure time, and particle diameter for thorough fusion of metal powder particles was laser corroborated principally by experimental observations. However, already in the past, sintering praxis has revealed that comprehensive fusion of the discrete particles will not suffice, neither for overall fusion of a sinter layer nor for the fusion of a new sinter layer with the substrate or the previous layers. This “excessive” energy demand occurs also in laser melting with cw-radiation. This is another finding of the investigations.OutlookRapid rate laser melting of bodies with highly resolved geometries can only be obtained with high scanning velocities.

This is another finding of the investigations.OutlookRapid rate laser melting of bodies with highly resolved geometries can only be obtained with high scanning velocities. As the fusion of the powder still requires the merging of the melt entities originating from the powder particles, the hydrodynamic properties or the limited fluidity of the melt might become a limiting factor for a technology that relies on short irradiation times. Therefore, effects that influence the melt dynamic e.g. by rapid boiling or plasma pressure, which have been observed in laser micro sintering with q-switched pulses, could eventually become necessary.

by rapid boiling or plasma pressure, which have been observed in laser micro sintering with q-switched pulses, could eventually become necessary. It will be an interesting question if cw-lasers, with significantly higher power than and with comparable beam quality as the presently available ones, will yield the chance to supplement the technology of rapid scanning laser melting with the melt condensing effects that have been observed with q-switched pulses.AppreciationsThe presented experiments have been funded by ‘Bundesministerium für Bildung und Forschung’ of Germany in the course of the project ‘INNOPROFILE – Transfer’. The authors gratefully acknowledge the preparation of the probes and the assembling of the setup by Christin Schulze and Martin Erler.[1][2][3][4][5][6][7][8]ReferencesC. Deckard, US patent 4,863,538. Filed: October 17th, 1986, published September 5th, 1989.Shellabear, M., Nyrhilä, O.. Proceedings 4th LANE 2004, Sept. 22.-24. 2004, M.

gh Resolution Topology Design with Iso-XFEMM. Abdi, I. Ashcroft, R. WildmanFaculty of Engineering, The University of Nottingham, Nottingham NG7 2RD, UKAbstractTopology optimization, as a challenging aspect of structural optimization, has gained interest in recent years as a method of designing structures to take advantage of the design freedoms of advanced manufacturing techniques such as Additive Manufacturing (AM). The majority of topology optimization algorithms are integrated with the Finite Element Method (FEM) to enable the analysis of structures with complex geometry during the optimization process. However, due to the finite element-based nature of the subsequent topology optimized solutions, the design boundaries are dependent on the finite element mesh used and tend not to have the desired smoothness for direct fabrication. The topology optimized solutions may, therefore, need smoothing, reanalysing and shape optimization before they become manufacturable.

The topology optimized solutions may, therefore, need smoothing, reanalysing and shape optimization before they become manufacturable. In this study, an Extended Finite Element Method (X-FEM) is employed and integrated with an evolutionary structural optimization algorithm, aiming to avoid/decrease the post-processing required from topology optimization design to manufacture. Rather than using finite elements for boundary representation, an isoline/isosurface approach is used to capture the design boundary during the optimization process. The comparison of the X-FEM-based solutions with the FE-based ones for the topology optimization of test cases representing real industrial components indicates significant improvements in the solutions’ boundary representation as well as their structural performance.IntroductionAdditive manufacturing (AM) has grown over the last 30 years to be used in various application areas.

It contrasts with the conventional fabrication processes which are usually subtractive or formative, in that the part is built up layer-by-layer. This layer manufacturing approach allows producing parts of significantly greater complexity compared with traditional processes and this increased complexity generally does not have a significant effect on the cost of the process. Not only this provides the designer with significantly greater design freedom, it also enables the built part to be closer to the optimum design than is possible with traditional processes (Brackett et al 2011).The level of freedom offered by AM processes necessitates the development of new design tools since the current design tools were developed prior to AM and are therefore better suited to traditional manufacturing methods. Topology optimization, however, offers great potential for AM, since it is capable of achieving solutions with great complexity.

Topology optimization, however, offers great potential for AM, since it is capable of achieving solutions with great complexity. Topology optimization is a structural optimization technique, mainly developed in the last three decades, which aims to find the best possible material distribution and connectivity for a structure within a specified design domain under a set of loads and boundary conditions. It involves finding the number of members required in the design and the manner in which these members areconnected. There have been a significant number of papers dedicated to this subject mainly based on homogenisation (Bendsøe and Kikuchi, 1988), Solid Isotropic Material with Penalization (SIMP) (Bendsøe, 1989; Zhou and Rozvany, 1991), Evolutionary Structural Optimization (ESO) (Xie and Steven 1993, 1997), level set method (Wang et al., 2003; Allaire et al, 2004) and genetic algorithms (Jakiela et al. 2000; Ryoo and Hajela 2004).

2000; Ryoo and Hajela 2004). Most of the proposed topology optimization algorithms have been applied to simple, standard problems, such as Michelle-type structures and cantilever beams with rectangular domains, there has been less attention on applying these algorithms to 3D real-life structures and loading scenarios. In some cases the mathematical complexity or the size of the FE design domain doesn’t allow the algorithm to be properly implemented. OptiStruct (Altair Engineering Inc.) is an example of software designed to enable the SIMP method of topology optimization to be applied to real components. Other software such as Nastran (MSC Software) and Abaqus FEA (Dassault Systèmes) also have options to apply similar density-based approaches to find the solution to topology optimization problems.

Other software such as Nastran (MSC Software) and Abaqus FEA (Dassault Systèmes) also have options to apply similar density-based approaches to find the solution to topology optimization problems. Although the topology optimization modules of these software applications are being widely used for research and engineering purposes, a drawback of the density-based approaches (and many other element-based approaches) is that they cannot provide a clear and smooth representation of the design boundaries in converged topologies. This issue brings difficulties in interpreting the solutions, combining them with CAD and manufacturing the topologies.

This issue brings difficulties in interpreting the solutions, combining them with CAD and manufacturing the topologies. Therefore the solutions usually need post-processing, reanalysing and shape optimization before manufacture, which is costly and time consuming.A possible alternative approach to the conventional element-based methods in topology optimization is the use of the eXtended Finite Element Method (X-FEM) combined with an implicit representation of the design boundaries (such as those in level set and isoline/isosurface approaches). X-FEM extends the classical finite element approach by adding special shape functions which can represent a discontinuity inside finite elements. Therefore in structural optimization applications, X-FEM can be used where the boundary of the design crosses finite elements, without the need to remesh the boundaries.

Therefore in structural optimization applications, X-FEM can be used where the boundary of the design crosses finite elements, without the need to remesh the boundaries. In the isoline approach (Maute and Ramm, 1995; Lee et al, 2007; Victoria et al, 2009) the boundary of a design is represented by isolines of a selected measure of structural performance behaviour, such as von Mises stress or strain energy density. In a previous study (Abdi et al 2014a), an evolutionary based optimization approach combined with X-FEM and isoline boundary representation was used for the topology optimization of 2D continuum domains. The numerical comparison of the converged solutions from Iso-XFEM approach with those obtained using bi-directional evolutionary structural optimization (BESO) showed the efficiency of the algorithm.

The numerical comparison of the converged solutions from Iso-XFEM approach with those obtained using bi-directional evolutionary structural optimization (BESO) showed the efficiency of the algorithm. The aim of this paper is to extend the Iso-XFEM method into 3D to address the application of X-FEM and isolines/isosurfaces of structural performance to the topology optimization of real-life structures. To meet this aim, 2D and 3D test cases including an industrial arm are considered and a comparison of the solutions with BESO (as a discrete element-based topology optimization method) solutions is presented.Isoline and Isosurface Boundary RepresentationIn general, isoline/isosurfaces are the lines/surfaces that represent the points of a constant value, named the isovalue, in a 2D/3D space. The isoline approach is an implicit method of defining the boundaries of a design in a 2D fixed grid design domain.

The isoline approach is an implicit method of defining the boundaries of a design in a 2D fixed grid design domain. In this approach the boundaries are defined using the contours of a higher dimension (3D) of structural performance,such as Strain Energy Density (SED) or von Mises stress, obtained from finite element analysis of the design space. The isosurface approach is the extension of the isoline method to the representation of the boundaries of a 3D design using contours of a 4D structural performance. In structural optimization applications (Victoria et al 2009, 2010; Abdi et al 2014a, 2014b) the boundaries are defined by the intersection of the structural performance (SP) distribution with a minimum level of performance (MLP), which is typically increasing during the optimization process.

Figure 1 shows a 2D fixed grid design domain discretized with a 30x30 mesh, where the intersection of the SED distribution as a structural performance criteria with a minimum level of SED gives the design boundary.The regions of the design domain in which the structural performance is lower than the MLP is defined as the void part of the domain and those which have a higher value than the MLP form the solid part of the design domain. The relative performance, , is defined as:(1)Following this, the design domain can be partitioned into void phase, boundary and solid phase, with respect to the values of relative performance, as( ) {( ) ( ) ( )(2)Figure 2 shows how the design space, D, from figure 1 is partitioned into , andusing the relative performance function ( ), distributed over the design space.Figure 1. Design boundary represented by intersection of a SP with the MLP.Figure 2. A 2D design domain represented by its relative structural performance.

Design boundary represented by intersection of a SP with the MLP.Figure 2. A 2D design domain represented by its relative structural performance. (a) Relative performance model. (b) Design domain.By implementing the above isoline/isosurface approach, the design boundary is superimposed on the fixed grid finite elements, making three groups of elements in the FE design space: solid elements, void elements, and boundary elements (the elements which lie on the boundary) as demonstrated in figure 3. The contribution of solid and void elements to the FE framework could simply be considered using a soft-kill scheme (Huang and Xie, 2009) in which instead of deleting the elements in the void phase, they are assigned a weak material property. However, in order to accurately represent the design boundary whilst avoiding expensive remeshing operations, an X-FEM approach can be employed as discussed in the next section.Figure 3.

However, in order to accurately represent the design boundary whilst avoiding expensive remeshing operations, an X-FEM approach can be employed as discussed in the next section.Figure 3. The elements are classified into 3 groups by superimposing the design geometry on the fixed grid FE mesh.X-FEM for Structural OptimizationX-FEM (Belytschko and Black, 1999; Moës et al, 1999) is a generalized approach for the classical finite element method which can represent discontinuities such as cracks, holes, inclusions, and fluid/structure interaction, inside finite elements without the need to remesh the internal boundaries. In the X-FEM, the classical finite element (FE) shape functions are extended by adding discontinuous shape functions to the displacement field in order to enrich the FE approximation space near the discontinuity. X-FEM has been applied to various kinds of discontinuity by defining an enrichment function appropriate to the nature of the discontinuity.

X-FEM has been applied to various kinds of discontinuity by defining an enrichment function appropriate to the nature of the discontinuity. In the case of structural optimization problems, X-FEM can be implemented to represent theevolving boundary of a design on a fixed grid FE mesh with no need for remeshing. In this case the elements which lie on the design boundary experience a material-void discontinuity. In this case, the X-FEM scheme for modelling holes and inclusions proposed by Sukumar et al (2001) can be utilized to represent the FE approximation space near the discontinuity:( ) ∑ ( )( ) (3)where are the classical FE shape functions, are the nodal degrees of freedom and H(x) is a Heaviside function with the following properties( ) {(4)implying that the value of the Heaviside function switches from 1 to 0 in the void parts of the design domain.

In order to realize the above X-FEM scheme, integration over the element domains are only performed on the solid sub-domain of the boundary elements while the fully solid/void elements are treated using the conventional FEM. Following this, the stiffness matrix of a 2D element can be defined by∫ ( )(5)where is the element domain, is the displacement differentiation matrix, is the elasticity matrix for the solid material and t is the thickness of the element.Various partitioning schemes can be implemented to realize the above X-FEM integration scheme, dependent to the element type that is used in the FE model of the domain of interest. In the case of 2D quad elements, we can partition the solid sub-domain of the elements into sub- triangles and use the Gauss quadrature method to numerically calculate the integral for each triangle. The element stiffness matrix can then be obtained by summing the contributions of each sub-triangle (figure 4).Figure 4.

The element stiffness matrix can then be obtained by summing the contributions of each sub-triangle (figure 4).Figure 4. Quadrilateral boundary elements represented by the 2D X-FEM scheme.Figure 5. Decomposing a boundary hexahedral element: (a) a boundary hexahedral element where the whole element domain is partitioned into 24 sub-tetrahedra; (b) a solid sub- tetrahedron; (c) a boundary sub-tetrahedron where the solid sub-domain is further partitioned into sub-tetrahedrons.In the case of 3D Hex elements, the solid sub-domain of the boundary elements are partitioned into sub-tetrahedrons and the integration is numerically performed over the solid tetrahedrons. In our implementation, the hexahedral boundary element is initially partitioned into 24 sub-tetrahedra by dividing each of the 6 surfaces of the element into 4 triangles and connecting the edges of the triangles to the centroid of the hexahedron. The void tetrahedra are then removed from the integration domain.

The void tetrahedra are then removed from the integration domain. The bi-material (solid/void) tetrahedra are again partitioned into smaller sub-tetrahedra and the void tetrahedra removed from the integration domain. Integration is eventually performed over the remaining solid tetrahedra (figure 5).Evolutionary based Optimization Algorithm using Isolines/IsosurfacesThe optimization algorithm developed in this study is evolutionary based, i.e. based on the simple assumption that the optimized solution can be achieved by gradually removing the inefficient material from the design domain. Examples of evolutionary based optimization methods include ESO in which inefficient elements are gradually removed throughout the optimization process, and BESO in which the elements are added and removed simultaneously throughout the process.

However, unlike ESO and BESO in which the material removal/addition is carried out at an elemental level, in this approach the optimization operates at a global level of structural performance by the use of isoline/isosurface design representation. An appropriate performance criterion is used to characterize the efficiency of material usage in the design domain. Material is then removed from low relative performance regions ( ( ) ) and redistributed to the high relative performance regions ( ( ) ). As shown in equation 1, the relative performance can be calculated by subtracting a minimum level of performance (MLP) from the structural performance (SP). However, in order to stabilize the evolutionaryprocess, as suggested by Huang and Xie (2007), historical information of the structural performance can be used by averaging the current relative performance with that from the previous iteration:Figure 6.

Flowchart of the optimization method.(6)where i denotes the node number and it is the current iteration number. The minimum level of performance (MLP) which is used in equation 1 to find the relative performance, is usually increasing during the evolutionary process and can be calculated from:/ (7)where is the maximum performance over the design domain and is the redistribution factor for the current iteration. is the current volume fraction of the solid material to the whole domain, and is used in the above equation to accelerate the material removal process at lower volume fractions. With the current redistribution factor, the iterative process of theextended finite element analysis and material removal/redistribution takes place until the percentage change in volume fraction is less than a minimum value , which means that a steady state is almost reached. Then, the redistribution factor is increased by adding an evolutionary rate, ER:.

Then, the redistribution factor is increased by adding an evolutionary rate, ER:. (8)With the new redistribution factor, the extended finite element analysis and material removal/redistribution is repeated until a new steady state is reached. The evolutionary process continues until a desired optimum, such as a final volume fraction ( ) is reached. Figure 6 summarizes the optimization algorithm used in this study.Test CasesIn order to illustrate the efficiency of the proposed Iso-XFEM algorithm, 2D and 3D test cases are presented in this section. A Matlab code was developed to apply the Iso-XFEM algorithm. The design domains of the test cases were more complex than the test cases used in the previous study (Abdi et al, 2013), and the mesh was not necessarily structured in these cases. Also, in the 3D test case, a non-design domain was considered which is usually the case in the topology optimization of real structures.

Also, in the 3D test case, a non-design domain was considered which is usually the case in the topology optimization of real structures. Therefore, due to the complexity of the initial geometries, the FE models of the structures were created in a FE package and then imported into Matlab in terms of Nodes and Elements matrices. All the results were generated using a desktop computer with an Intel Xeon 2 processor of 2.4 GHz speed and 24 GB RAM.Figure 7. Design domain, FE mesh, loads and boundary conditions used for the C clip.Figure 8. Evolution of the topology of the C clip.Test case 1: a two dimensional C clipIn this example, the Iso-XFEM approach was used to optimize a C clip with the loads and boundary conditions shown in figure 7. Node A located on the middle of the left side of the clip is fixed in the x direction and node B on the right is fixed in both x and y directions.

Node A located on the middle of the left side of the clip is fixed in the x direction and node B on the right is fixed in both x and y directions. Here, the optimization problem was defined as minimizing the strain energy of the clip subject to a final volume fraction, , and the strain energy density of the design domain is used as the structural performance criterion. The material used had a Young’s modulus of 210 GPa and a Poisson’s ratio of 0.3. The optimization parameters used for this test case were and ∆V = 0.005. Figure 8 shows the resulting topologies for a range of volume fractions. It can beseen that the process of material removal was carried out by creating a number of holes in the low SED regions of the design space, followed by the merging and enlarging of holes until the desired volume fraction and convergence conditions were achieved. The triangles near the boundary represent the X-FEM integration domains and are not elements.

The triangles near the boundary represent the X-FEM integration domains and are not elements. Therefore as no element is added or removed to/from the design domain, the number of degrees of freedom during the optimization process hasn’t been changed. However, the boundary of the converged solution (figure 9) is smoothly defined by implementing the isoline and X-FEM approaches. The optimization converged after 120 evolutionary iterations which took 170 seconds.Figure 9. Converged solution for VF = 0.4 at iteration 120.Test case 2: an industrial armIn order to implement the Iso-XFEM topology optimization method on a practical problem, the industrial arm shown in figure 10 which was previously studied using BESO (Aremu et al, 2013), was considered. A load of 667.23 N was uniformly distributed on the lower edge of surface A, and all degrees of freedom on surface B were fixed, forming a cantilever beam.

A load of 667.23 N was uniformly distributed on the lower edge of surface A, and all degrees of freedom on surface B were fixed, forming a cantilever beam. Two levels of difficulty existed in this problem compared to the 2D/3D cantilever problems often used in the literature; an increasing geometrical complexity and existence of non- design domain. The grey cylindrical regions in figure 10 were set as non-design domain and the rest of the structure was the design domain. The objective was to minimize the total strain energy subject to a final volume fraction of 15% of the design domain. The material properties of the arm were a Young’s modulus E =74 GPa and Poisson’s ratio υ = 0.33. Two experiments employing different mesh sizes were considered for this test case. In each of the experiments, the optimized solution was obtained using the proposed Iso-XFEM approach and the element-based BESO method.

In each of the experiments, the optimized solution was obtained using the proposed Iso-XFEM approach and the element-based BESO method. This was to enable the effectiveness and efficiency of the two approaches to be compared.Experiment 1The whole domain was meshed using approximately 22000 hexahedral elements. The optimization parameters were and . The resulting topologies obtained using the Iso-XFEM method for a range of volume fractions up to 0.15 are shown in figure 11. Employing the same mesh used for the previous experiment, the topology optimization problemof the arm was solved using a soft-kill BESO code based on Huang and Xie’s study (2007). The BESO parameters used were an evolution rate, ER = 0.02, and sensitivity filter radius of 1.2 times the average element size. The converged solutions of the two approaches are shown in figure 12.Figure 10. Design domain and non-design domain of the industrial arm.Figure 11.

The converged solutions of the two approaches are shown in figure 12.Figure 10. Design domain and non-design domain of the industrial arm.Figure 11. Evolution of the topology of the industrial arm.Figure 12. Converged solutions of experiment 1: (a) Iso-XFEM solution and (b) BESO solution for VF = 0.15. Surface roughness was measured in regions Z1 and Z2.Figure 13. Converged solutions in experiment 2: (a) Iso-XFEM solution and (b) BESO solution for VF = 0.15.Experiment 2In this experiment, the whole domain of the industrial arm was discretised using approximately 32000 Hex elements. The topology optimization of the arm was considered using both the BESO and Iso-XFEM optimization approaches with the same optimization parameters as were used in experiment 1. Figure 13 shows the converged solutions obtained using the two different approaches.

Figure 13 shows the converged solutions obtained using the two different approaches. It can be seen that the converged solutions in experiment 2 have similar topologies to those seen in experiment 1 (figure 12), apart from the fact that topologies with more members have been realized by using the finer mesh in experiment 2 (table 1).DiscussionConsidering experiment 1, it can be seen that both Iso-X-FEM and BESO approaches have successfully obtained converged solutions for the topology optimization problem of the industrial arm as an example of a real structure. It can be seen that the overall topologies of the converged solutions look similar; however the Iso- XFEM solution has a much smoother surface than the BESO solution. Measuring the surface roughness of the solutions is not very easy as it varies depending on the geometrical features forming the topology.

Measuring the surface roughness of the solutions is not very easy as it varies depending on the geometrical features forming the topology. Assuming that spherical and cylindrical features have higher values of surface roughness than flat features, two regions Z1 (hemispherical) and Z2 (cylindrical) in figure 12 for experiment 1 (and similar regions in figure 13 for experiment 2) were considered and their arithmetic mean surface roughness, , was calculated, as shown in table 1. It can be seen that in both regions, much higher values of surface roughness are obtained with the BESO solution. The jagged edges of the coarse finite elements on the boundary of the BESO solution (figure 12b) has resulted in an unfeasible design boundary which needs further interpretation of the topology and post processing such as smoothing, reanalysing and shape optimization before manufacture.

Comparing the strain energies of the BESO and Iso-XFEM solutions in table 1, it can be seen that the Iso-XFEM approach has resulted in a performance for the converged solution approximately 18% higher than the BESO solution when compared with the same mesh size. It can also be seen that the BESO solution has fewer members (less complexity) than the Iso-XFEM solution (tale 1). A reason for this could be the sensitivity filtering scheme (Sigmond and Peterson, 1998) that is employed with the BESO method (and many other element-based methods).

A reason for this could be the sensitivity filtering scheme (Sigmond and Peterson, 1998) that is employed with the BESO method (and many other element-based methods). This scheme is used to eliminate checkerboard pattern problems, however this filtration scheme is not required for the proposed Iso-XFEM method, although it can be used if solutions with less complexity are desired.Comparing the time cost of the two optimization approaches in experiment 1 (table 1), one may notice that the Iso-XFEM optimization approach operates slower than the BESO method for the same initial mesh as it takes more time to calculate the properties of boundary elements using the X-FEM integration scheme. However the time spent to solve the finite element linear system of equations will be approximately the same, as the same number of degrees of freedom exists in the both methods.

However the time spent to solve the finite element linear system of equations will be approximately the same, as the same number of degrees of freedom exists in the both methods. Therefore it is expected that by increasing the mesh density which results in increasing the number of degrees of freedom, the percentage difference between the solution time of the two approaches decreases. This issue was investigated by performing experiment 2 in which the design domain is discretised with a finer mesh than experiment 1.Table 1. Comparison of the solutions in experiment 1 and experiment 2. | | 45% 66% 97% 4.2% 9% 30% 27% 49% 32% 20% 6%Exp. I 22000 155170 78636 1659 1945 20 15 0.037 0.62 0.15 0.65Exp.

Comparing the effect of mesh refinement on the optimization’s time cost and the solutions’ objective of the two approaches, one can see that the Iso-XFEM optimization approach is less dependent on the mesh refinement as it has lower percentages of difference of time cost and objective in the two experiments. Also the comparison shows that when a finer mesh is employed, the time cost of the Iso-XFEM optimization approach becomes closer to the BESO optimization approach. It can be seen that the time cost of BESO in experiment 2 is almost equal to the time cost of Iso-XFEM in experiment 1 (approximately 155 ks). Using the values obtained for this time cost, table 2 compares the objectives and surface roughness of Iso-XFEM and BESO solutions for the same solution time. From table 2 it can be seen that even employing lower number of elements for the Iso-XFEM approach, the objective is still lower than the BESO one, showing higher performance of the Iso-XFEM solutions at the same time cost.

Again for the same solution time, it can be seen that the Iso-XFEM solutions are represented with more smooth boundaries than the BESO solutions.ConclusionsIn this paper, the Iso-XFEM approach is developed and considered for generating high resolution topology optimization solutions for real-life structures including a 3D industrial arm. It is shown that the combination of the X-FEM and isoline/isosurface boundary representation can be effectively used to generate efficient and smooth topologies from complex design domains. Although the use of X-FEM in structural optimization of continuum structures can increase the solution time, it is shown that this is offset by the benefit of being able to employ acoarse mesh to generate a smooth topology.

gh Speed Sintering - Continuing research into a new RapidManufacturing process Helen R Thomas, Neil Hopkinson, and Poonjolai Erasenthiran Rapid Manufacturing Research Group, Loughborough University, UKAbstractHigh Speed Sintering (HSS) is an emerging layer manufacturing technique aiming to break into the lucrative field of Rapid Manufacturing (RM). The process is likened to Selective Laser Sintering (SLS), however, instead of a laser dictating the sintered cross sectional area of each layer, the desired area is first printed using a Radiation Absorbing Material (RAM) and then sintered using an inexpensive infrared lamp. This paper begins by describing the sintering process in more detail and then outlining the overall manufacturing cycle. It then continues by describing the experiments performed to investigate the current problem concerning the hardness of excess powder within the powder bed.

It then continues by describing the experiments performed to investigate the current problem concerning the hardness of excess powder within the powder bed. This problem arose due to the continual exposure of the whole bed to infrared radiation from the lamp. The experiments showed that as the power of the IR lamp increased, the hardness of the bed also increased. Furthermore, at higher IR power levels it was found the excess powder produced a solid tile which could only be broken down by a glass bead blaster.KeywordsRapid Manufacturing, High Speed Sintering, Inkjet Printing, Infrared Radiation.Rapid ManufacturingRapid manufacturing, (the production of end use products by layer manufacturing) is becoming a viable manufacturing process within numerous engineering firms. Existing studies [1,2] have shown that selective laser sintering can be used to manufacture small components in volumes up to 14,000.

Existing studies [1,2] have shown that selective laser sintering can be used to manufacture small components in volumes up to 14,000. After this, machine cost and build speed inhibit any higher economical production runs. Furthermore, Phonak Hearing Systems and Siemens Hearing Instruments have drawn on Stereolithography’s qualities by using the technology to manufacture bespoke hearing aids. With SL, a batch of 75-100 hearing aids shells can be manufactured in six to eight hours [3].In recent years, a number of other organizations have begun to pursue the design of layer manufacturing technologies which directly cater for Rapid Manufacturing. Speedpart [4] have developed and marketed a process that masks a plastic powder bed which is then exposed to IR radiation. This enables sintering of an entire layer in a layer manufacturing cycle time of 12-14 seconds [3].

This enables sintering of an entire layer in a layer manufacturing cycle time of 12-14 seconds [3]. Desktop Factory [5] also have a machine on the market which uses an inexpensive halogen light source and drum printing technology to build parts layer by layer from plastic powder [3]. In comparison to SLS and HSS, their build area is small; however their manufacturing method is similar to the technologies discussed here. Finally, two university based research projects are also developing potential RM processes. Firstly, the University of Southern California is developing a process called Selective Inhibition Sintering (SIS) [6,7]. This process prints a sintering inhibitor at the outline of a desired geometry. Infrared radiation then exposes each layer, sintering the desired cross- sectional area. Finally, electrophotographic rapid prototyping (ERP) is being developed at the University of Florida. The process works by picking up and depositing powder using electrostatic forces [8,9].

Finally, electrophotographic rapid prototyping (ERP) is being developed at the University of Florida. The process works by picking up and depositing powder using electrostatic forces [8,9]. Negative charge is deposited onto a photoconductor drum. A UV682beam then scans and discharges selective regions upon the surface of the drum. Powder is deposited on these newly created discharged areas and as the roller rotates, then transferred to the build area. Each layer is sintered by exposure to radiant heat [8,9].These processes possess varying manufacturing methods, yet all have two key aspects which provide them with the potential to be RM processes; firstly, sintering occurs in a time that is irrespective of the size, amount and shape of the 2D profiles in each layer and secondly, the deposition or masking of a 2D profile is performed relatively quickly.

These advantageous characteristics enable part cost to be reduced and therefore open up opportunities to manufacture at increasing part volumes.Introduction to High Speed SinteringIn the last three years, a new layer manufacturing process has been developed at Loughborough University. The process has some similarities to both Selective Laser Sintering (SLS) and 3DPrinting (3DP); however its novel differences suggest that in the future it may possess the capability to compete with injection moulding. Instead of an expensive laser dictating the sintered cross sectional area of each layer, the desired area is first printed using a radiation absorbing material (RAM) and then sintered by exposing the whole part bed to infrared (IR) radiation – see Figure 1. The key to the process is that the RAM absorbs the IR radiation at a higher rate than the unprinted areas of Nylon powder.

The key to the process is that the RAM absorbs the IR radiation at a higher rate than the unprinted areas of Nylon powder. In a matter of seconds, the printed RAM absorbs and transfers enough thermal energy to enable the underlying nylon powder to reach its melt temperature and initiate viscous sintering. This process has been named ‘High Speed Sintering’ (HSS) and international patent applications have been applied for.Infrared radiationSintered area of printed RAMInkjet printheadUnsintered excess Nylon powderFigure 1 – High Speed SinteringHSS offers many advantages over SLS. The removal of the laser within HSS not only reduces the machine cost significantly but also increases the build speed dramatically. The increased build speed is a result of the infrared radiation system; the time consuming scanning of the laser across the powder bed is eliminated and sintering occurs in a time that is irrespective of the size, amount and shape of the 2D profiles in each layer.

These advantageous characteristics hold the key to HSS’s progression into the field of rapid manufacturing.Process DevelopmentIn 2003 Hopkinson and Erasenthiran proved that the addition of carbon black to a standard Nylon powder increased the rate of sintering such that an entire layer may be sintered in 5 seconds using an infra-red lamp [10]. They then manufactured tensile test683specimens via HSS and found that the mechanical properties of the HSS material were better than that of equivalent SLS parts – see Table 1 [10].ProcessSLS*% (by weight ) of carbon black mixed with Duraform 0Youngs’ Modulus (MPa)1600UTS (MPa)44Elongation at break (%)9HSS0.25163347.518HSS2166646.415Table 1 - HSS and SLS part propertiesprovided by 3Dsystems.comThe success of the initial experiments performed by Hopkinson and Erasenthiran led to industrial support of the project.

An SLS machine was given from 3D systems to the Rapid Manufacturing Research Group (RMRG) at Loughborough University enabling the creation of a HSS test rig. The test rig was then used to experiment HSS via masking and to begin investigating HSS via deposition of a secondary material to promote energy absorbance.HSS via masking was investigated by Thomas, H.R. in 2005. A series of masks were attached to the roller and the bed was filled with premixed Nylon 12 and carbon black (2% by weight). The unmasked area of the bed was then exposed to IR radiation. Sintering was successful; although, the repeatability of the roller position was poor. This made it difficult to produce multi layer parts as it was impossible to place the mask in the same position each time.

This made it difficult to produce multi layer parts as it was impossible to place the mask in the same position each time. At this point it was decided to proceed onto the second method of sintering.The second method (HSS via deposition of a secondary material to promote energy absorbance) described was achieved by selectively printing a Radiation Absorbing Material (RAM) onto the powder bed. The HSS test rig was adapted for this new role by attaching an inkjet printhead and Infrared lamp onto the roller. This setup is illustrated in Figure 2 and Figure 3.Figure 2 – The HSS machineA HSS manufacturing method was then developed and automated by creating and running a C programme to control the position of the roller, powder bed heights, when the684inkjet printhead prints and when the IR lamp exposes the bed to IR radiation.

Figure 3 illustrates a close up of the build area.Figure 3 – A close up of the HSS build areaThe manufacturing method is currently being optimized; however, the current cycleoperates as follows:o A layer of Nylon powder is deposited onto the build area (from left to right). o Using a radiation absorbing material (RAM), a 2D profile is then printed onto theNylon powdero Following the printhead, an IR element, exposes the image as the roller traversesthe bed (from right to left).o The RAM then absorbs the IR radiation at a higher rate than the Nylon powder. o The RAMs thermal energy is then transferred to the nearby Nylon powder. The Nylon powder then reaches its melt temperature and viscous sintering is initiated.o The desired 2D profile has now been sintered.

The Nylon powder then reaches its melt temperature and viscous sintering is initiated.o The desired 2D profile has now been sintered. o Another layer of powder is deposited onto the build area and the process isrepeated until the required geometry has been manufactured.Current Manufacturing IssuesOne of the main issues that arose during the process development of HSS was the hardness of the excess powder within the HSS part bed. Compared to SLS, the HSS part bed receives more thermal energy due to the scanning of the whole part bed by the Infrared lamp thus causing partial sintering. Although the partial sintering of the excess powder does not affect the properties of the part, it does affect part removal. Therefore, it was decided to explore this problem further and investigate the effect of a change in IR power level, upon the hardness of excess powder within the HSS part bed.

Therefore, it was decided to explore this problem further and investigate the effect of a change in IR power level, upon the hardness of excess powder within the HSS part bed. It was anticipated that the results from the investigation would highlight which heat sources provided the thermal energy that caused685the partial sintering of the main part bed. This information would then be used to eradicate the problem and enable easy excess powder removal within the process.Initial observations of the process had shown that once a part has been manufactured, and the machine allowed to cool down overnight, more than half the excess powder within the part bed had been partially sintered. It was also found that the excess powder exhibited different characteristics, depending upon its position within the bed.

It was also found that the excess powder exhibited different characteristics, depending upon its position within the bed. Figure 4 illustrates three areas in which distinct differences were observed.Figure 4 – Areas of partial sintering within the HSS part bedPowder in area 1 remained soft and was very easily removed from the bed. Powder in area two was significantly harder and a scalpel had to be used to break away the powder. Finally, powder in area three was almost as hard as the part itself and had to be removed using a shot blaster – see Figure 5.Figure 5 – Partially sintered excess powder around partDue to the variations in powder hardness across the whole part bed it was decided to perform a series of experiments that would map the hardness of the excess powder throughout the whole powder bed.686The experiments were performed by processing 20mm of Nylon 12 powder using different part bed heater temperatures and IR power levels – see Table 2.

Additionally, no RAM was printed onto the bed during the experiments. As a result, this eliminated any effect that the thermal energy emitted from a part would have on the excess powder within the bed.Experiment Number 1a, 1b, 1c 2a, 2b, 2cTemperature setting for IR part bed Heater 100 105IR lamp Power level ( 0 – 10 ) 0, 5, 7 0, 5, 7Table 2 - Experimental IR power levels to test powder hardnessFinally, it is important to note that, at present, the temperatures of the part bed heaters are not calibrated to °C and the numbers merely provide a reference point for the experiments.MethodologyEach experiment started by placing virgin powder into the left hand side feed bed andmain part bed. The initial levels of the beds are shown in Figure 6.125mm20mmOverflow powderFigure 6 – Powder bed heights (Powder Hardness)The machine was then left to preheat for two hours.

The initial levels of the beds are shown in Figure 6.125mm20mmOverflow powderFigure 6 – Powder bed heights (Powder Hardness)The machine was then left to preheat for two hours. Once the main part bed and left hand side feed bed had preheated to appropriate temperatures, the beds were then raised to their manufacturing temperatures. The C programme was then programmed to process 200, 0.1mm, layers and the build was started. Once the cycle had finished, the bed heaters were all switched off and the machine was left to cool down overnight. The following morning, the build chamber was opened and a mask placed onto the powder bed – see Figure 7.Figure 7 – Plate to locate positions to test powder687A random number generator was then used to create a list of the order in which each hole was to be tested.

A Ridsdale hardness tester (normally used to test the hardness of compacted sand in a casting mould) was then inserted into each hole and a hardness level was recorded – see Figure 8.A tested hole within the part bedRidsdale hardness testerMask over the powder bedFigure 8 – Taking a powder hardness readingOnce all the measurements had been taken, the powder was removed from the machine and fresh virgin powder placed into the machine. The next experimental variables were then tested in exactly the same way.One further experiment that was performed was the measurement of bed hardness in an SLS machine. The measurements of the powder bed hardness were taken after a normal SLS build and recorded in exactly the same way that is described above.

The measurements of the powder bed hardness were taken after a normal SLS build and recorded in exactly the same way that is described above. The results are shown in the next section and are used as a comparison to the hardness of the HSS bed.Results and Discussion10090807060 5040 30 20 10 0S3S7S51234567S1Figure 9 – The SLS part bed, part bed heater = 172˚C688Figure 9 illustrates the powder bed hardness in an SLS machine. The part bed contains soft powder and the SLS parts were easily removed from the bed. The results demonstrate that in this case, the SLS part bed heater had very little effect upon the hardness of the excess powder. Furthermore, it was observed that the excess powder located next to the SLS parts exhibited a similar softness to that of the rest of the powder bed.

This leads to the observational conclusion that SLS parts do not emit enough thermal energy to the surrounding excess material that initiates any degree of sintering.The next three figures illustrate the first set of HSS part bed harness tests.100 90 80 70 60 50 40 30 20 10 0S7S5S31234567S1Figure 10 – Exp 1a: Part bed = 100, Infrared power level = 0100100 9080 70 60 50 40 30 20 10 0S7S5S39080 70 60 50 40 30 20 10 01234567S11234567Figure 11 - Exp 1b: Part bed = 100, Infrared power level = 5Figure 12 - Exp 1c: Part bed = 100, Infrared power level = 7This first set of HSS results illustrate the range of powder bed hardnesses when thepowder bed heater was set to 100 and the IR power; 0, 5 and 7.689S1S7S5S3Figure 10 shows that when the bed is not exposed to IR radiation from the lamp, itremains soft. This is a similar powder hardness to that achieved in SLS – see Figure 9.

This is a similar powder hardness to that achieved in SLS – see Figure 9. However, once the bed is exposed to IR power levels 5 and 7, the hardness increases dramatically to level that would make it very difficult to remove parts. At power level 5, the stiff tile can be broken by hand however, at power level 7, the powder tile can only be broken down by a scalpel or shot blaster.

Furthermore, the results show that the immediate front and back of the machine remain soft due to firstly, the IR radiation emitted from the lamp not exposing the extremes of the bed and secondly, the thermal energy within the powder bed is able to dissipate through the machine bed walls.The next three images show the results taken when the part bed heater was set to 105.10090 80 70 60 50 40 30 20 10 0S7S5S31234567S1Figure 13 – Exp 2a: Part bed = 105, Infrared power level = 010090 80 70 60 50 40 30 20 10 0S7S5100 9080 70 60 50 40 30 20 10 0S31234567S11234567Figure 14 - Exp 2b: Part bed = 105, Infrared power level = 5Figure 15 - Exp 2c: Part bed = 105, Infrared power level = 7The second set of results illustrates a further increase in powder bed hardness when the part bed is exposed to IR radiation.

To begin with, when the bed is only exposed to690S1S7S5S3thermal radiation from the part bed heater, it exhibits a medium powder hardness - Figure 13. At this level, the powder can be broken up by applying some force with hands. Compared toFigure 10, there is a significant difference, thus providing the conclusion that the part bed heater does have an effect upon the part bed hardness within the HSS test rig.Figure 14 and 15 illustrate further the effect of exposing the already medium level powder hardness to IR radiation from the lamp. In these two experiments, the tile was solid and could only be broken up using a scalpel or shot blaster.Experimental ConclusionsThe experiments have shown that the additional thermal energy delivered to the bed by the IR lamp, results in an increased hardness of the powder. As the power of the IR lamp increases, the hardness of the bed also increases.

As the power of the IR lamp increases, the hardness of the bed also increases. However, the bed hardness is also influenced by the power of the main part bed heater. This illustrates the necessity to balance the power of part bed heater and IR lamp to enable easy powder removal. To understand the problem further, the following studies are planned:1.Investigation to determine what kind of sintering is occurring to the excess powder within the main part bed i.e. solid state sintering or viscous sintering.2. Measurement of the amount of thermal energy emitted from the IR lamp at power levels 0–10.3. Calibration and measurement of the amount of thermal energy emitted from the part bed heater.4. Measurement of the absorbency of Nylon 12 at a range of IR wavelengths.References1. Hopkinson, N. and Dickens, P.M. (2001), “Rapid prototyping for direct manufacture”, Rapid Prototyping Journal, Vol 7 No.4, pp.197-202.2. Hopkinson, N. and Dickens, P.M.

gh Speed Sintering – Early Research into a New Rapid Manufacturing ProcessAuthors:Neil Hopkinson & Poonjolai Erasenthiran Rapid Manufacturing Research Group Department of Mechanical and Manufacturing Engineering Loughborough University UKAbstract Rapid Manufacturing (the production of end use products by layer manufacturing techniques) has grown significantly in recent years and has started to revolutionise some areas of manufacturing. Among the main drawbacks for commercially available techniques are machine cost and build speed. This paper describes some initial research into a new process called High Speed Sintering. The High Speed Sintering process (UK patent No. 0317387.9) involves the sintering of 2D profiles of layers of powder without the need for a laser.

The High Speed Sintering process (UK patent No. 0317387.9) involves the sintering of 2D profiles of layers of powder without the need for a laser. Experiments performed on a simple lab apparatus have shown how the addition of carbon black to standard nylon powder can increase the rate of sintering such that an entire layer may be sintered in 5 seconds using an infra-red lamp. The effects of composition of carbon black on material properties are shown and may be traded off against build speed. Thermal control of the process is vital and the effects of altering the position and power used with an infra-red lamp are presented. Eliminating a laser reduces machine cost and build time, combining these factors will make the High Speed Sintering process suitable for high volume manufacture.

Eliminating a laser reduces machine cost and build time, combining these factors will make the High Speed Sintering process suitable for high volume manufacture. Cost predictions show that the process will be viable for the manufacture of standard products in volumes over 100,000.Background A previous study by the Rapid Manufacturing Research Group (then at De Montfort University, UK) in 2000 into the economics of using RP machines for medium to high volume manufacture suggested that selective laser sintering (SLS) machines may be used to produce small components in volumes up to 14,000 more economically than injection moulding (1,2). Since that time SLS has evolved as the most suitable process for Rapid Manufacture (RM) among the current range of available technologies and is used to manufacture products as diverse as bespoke hearing aids, aircraft ducts and parts for formula 1 racing cars (3).

Despite these successes, SLS has not been used for medium to high volume series production (hearing aids are produced in high volumes but each one is bespoke so this does not constitute high volume series production). The economic study in 2000 indicated that one of the major sources of cost for parts made in series production on SLS machines was machine cost. Machine cost for parts is dictated by the cost of the equipment required for manufacture and the speed of production achieved. Thus by reducing the cost of a layer manufacturing machine and increasing its throughput, a far more competitive process, suitable for medium to high volume series production may be achieved. The most obvious way to reduce the cost and increase the speed of an SLS machine is to eliminate the requirement for a laser. Eliminating the laser requires an alternative method for selectively sintering powder.

Eliminating the laser requires an alternative method for selectively sintering powder. The High Speed Sintering (HSS) process works by varying the absorbance of incoming radiant energy such that selected areas312of the surface absorb sufficient energy to elevate the temperature to the melt point while other areas, that absorb less energy, do not exceed the melt point (see Figure 1).Radiated energyLoose un- sintered powderSintered powderLayer 2 Layer 1Figure 1. Cross sectional view of the High Speed Sintering ProcessThere are a number of methods that will allow selected sections of the surface of the build bed to absorb more radiation than other areas including:1.

2.3.Using a part powder that absorbs energy more readily than the support powder The addition of a secondary material to promote energy absorbance in the required areas Using a mask to restrict incoming energy to selected areas of the surfaceThe work reported here involved the addition of different amounts of carbon black to nylon powder (as suggested in approach 2 above) but also used a mask to selectively sinter the shapes required (as suggested in approach 3 above). The use of a mask as suggested in approach 3 is effectively the method used in the selective mask sintering process that is being developed by Speedpart (4). Rather than using a mask it would be possible to deposit the secondary material as suggested in approach 2, for example by using conventional jetting techniques. Printing of a secondary material is an attractive option as it is a widely used technology that is capable of achieving resolutions tighter than those of current SLS.

Printing of a secondary material is an attractive option as it is a widely used technology that is capable of achieving resolutions tighter than those of current SLS. However in this initial experimental work the secondary material was simply pre-mixed with nylon powder and masks employed to achieve selective sinteringExperimental Method Powder mixing was performed using Duraform Nylon 12 powder along with carbon black MONARCH® 800 by first weighing the powder and then mixing with 10 KG PA material by rolling the tub for 2 hours. Figure 2 shows a simple bench top apparatus that was used to sinter various combinations of nylon/carbon black powder using a 2KW short wave infra-red lamp.313Infra-red lampPowder bedFigure 2. Bench top High Speed Sintering apparatusPrior to performing sintering experiments a thermal imaging camera was used to observe the temperature variations across a small exposed layer of powder.

Bench top High Speed Sintering apparatusPrior to performing sintering experiments a thermal imaging camera was used to observe the temperature variations across a small exposed layer of powder. Having established a reasonably even temperature distribution across the powder surface, experiments were performed to investigate the time required to sinter different combinations of powder from room temperature. Along with the sintering rate, observations of shrinkage rate and powder transfer during sintering were made. The bench top apparatus had no method for pre-heating, this resulted in warpage of samples and made the build up of multi-layer components very difficult. In order to overcome this, a Vanguard SLS machine from 3D Systems was modified so that sintering via the infra-red lamp could be achieved in more conventionally pre-heated SLS conditions. Figure 3 shows the modified Vanguard machine with the infra-red lamp secured to the roller mechanism.

Figure 3 shows the modified Vanguard machine with the infra-red lamp secured to the roller mechanism. The modified SLS machine allowed for the production of multi-layered parts in a variety of powder compositions.IR lampPart bedrollermaskFigure 3. Infra-red lamp and mask mounted to the roller mechanism on an SLS machine314One problem with the SLS machine was the repeatability of the position of the roller mechanism and hence mask on subsequent layers. Figure 4 shows the result of this lack of repeatability in positioning between layers.xFigure 4.75mmFigure 4. Tensile test piece created on the Vanguard SLS machine using a mask with a tensile test shapeRather than creating a mask in the shape of a tensile test specimen, a series of rectangular slabs of the required height were built and these were then cut using a CNC milling machine to create tensile test specimens that conform to BS EN ISO 527-2:1996.

All samples were cut with a smooth finish to avoid any surface cracks that may invalidate tensile test results. Figure 5a shows a rectangular slab comprising 20 x 0.1mm layers of nylon powder with added carbon black. Figure 5b shows a tensile test piece cut by CNC from a rectangular slab shown in Figure 5a.Figure 5a. Rectangular slab from the SLS Vanguard machineFigure 5b. Tensile test piece machinedfrom a rectangular slab.Parts made on the initial desk top system were subject to SEM analysis in order to view the nature of sintering and porosity in sintered samples.Results Figure 6 shows the temperature gradient across the surface of a small (70mm x 50mm) powder bed used in the initial tests for desk top sintering. Figure 6a shows the temperature gradient when the IR lamp was suspended 40mm above the bed surface.

Figure 6a shows the temperature gradient when the IR lamp was suspended 40mm above the bed surface. Figure 6b shows a considerable improvement in the temperature gradient achieved when the IR lamp was suspended at a height of 70mm above the bed surface.315Figure 6a. IR lamp 40mm above surfaceFigure 6b. IR lamp 70mm above surfaceFigure 6. Temperature variation across the powder bed surface for desktop sinteringFigure 7 shows the effects of percentage composition of carbon black (by weight) mixed with nylon powder on the time required to sinter powder. As expected carbon black absorbs infra-red energy at a higher rate than nylon and heats up more quickly than nylon. It is suspected that the particles of carbon black heat up and transfer thermal energy to surrounding particles of nylon by conduction and radiation thus causing the nylon to reach its melt temperature and sinter.

Figure 7 also shows that increasing the intensity of infra-red radiation increases the rate of sintering.3025Low Intensity) c e S20( e m T g n i r e t n Sii1510High Intensity5000.10.20.30.40.50.60.70.80.911.1Carbon Black (%)Figure 7. Effects of carbon black and radiation intensity on sintering time for nylonFigure 8 shows the degree of shrinkage caused during sintering through a square mask measuring 20mmx20mm. The size of the sintered material inside the unmasked area is ~16mmx16mm representing a significant linear shrinkage of ~20%.31620mmAluminium maskSintered powder creating solid part with shrinkageUn-sintered powderFigure 8. Shrinkage caused during sintering through a maskTable 1 shows results from tensile tests of parts tested according to BS EN ISO 527- 2:1996. The results show that parts created by HSS using carbon black have superior tensile properties to conventionally sintered SLS parts.

The results show that parts created by HSS using carbon black have superior tensile properties to conventionally sintered SLS parts. Of particular note is the increase in elongation to break of HSS parts as this property represents a fallibility for SLS in some end use applications.ProcessComposition of carbon black (% by weight) 0 0.25 2Young’s Modulus (MPa)SLS* HSS HSS * details obtained from www.3dsystems.com1600 1633 1666UTS (MPa)44 47.5 46.4Elongation at break (%)9 18 15Table 1. Effects of carbon black on mechanical properties of parts made by HSSFigure 9 shows SEM images from the surface of single layer parts produced on the desktop sintering system with magnifications of 50x (a) and 250x (b). The images show a high degree of necking between nylon particles suggesting that good mechanical performance should be achieved (this is consistent with the tensile test results presented in Table 1).

To date no similar images for parts created on the Vanguard machine under pre-heated conditions have been taken. It is possible that the parts produced on the Vanguard machines have a higher degree of necking due to pre-heating. The images in Figure 9 show small light coloured particles, it s not clear yet if these are carbon particles (D50 ~ 2µm) that are present on the surface of the larger nylon particles (D50 ~ 60 µm), EDX analysis will be required to establish this.317(a)(b)Figure 9. SEM images of sintered partsPotential for high volume manufacture In order to compare machine costs per part for HSS with those for SLS, the results from a previous study (1,2) have been used. Figure 10 shows a plan view of how a series of the parts used in the previous study could be oriented on a platform the same size as that used in the previous study (EOSP360), this is a layer of 70 parts.Figure 10.

Plan view of layout of 70 parts on a build bedEach part in Figure 10 has a height of 12mm (=120 slices) so an allowance of 150 slices for each layer of parts is assumed. In order to build 1050 parts, 15 layers of parts would be required with a total of 150x15 = 2250 slices. Figure 7 shows that sintering of a complete layer can be completed in 5 seconds, so allowing time for powder deposition (5 seconds), pre-heating (5 seconds) and depositing a radiation absorbing material (15 seconds) then a total time per slice of 30 seconds may be assumed. This results in a total build time of 1125 minutes (18.75 hours) to build 1050 parts. Table 2 shows the details from the previous cost analysis along with assumed data for an HSS machine with calculated machine costs to produce parts shown in Figure 10.

Machine costs from previous analysis with SLS and potential using HSSDiscussion The cost analysis shown in Table 2 assumes that HSS will use a build platform with similar dimensions to a current SLS machine (EOSP360). However, for higher volume production, platforms with a larger area will be able to increase throughput significantly and hence reduce machine costs further. The use of a secondary radiation absorbing material (in this case carbon black) is likely to add some cost to materials when compared with SLS. However increased use of powder materials for layer wise sintering processes are likely to bring down costs by economies of scale. The figures for SLS shown in Table 2 suggested that SLS may be used to manufacture products in volumes up to 14,000 at a lower cost than by injection moulding (1,2). The figures for HSS in Table 2 suggest that this process should be economically than 14,000.

The figures for HSS in Table 2 suggest that this process should be economically than 14,000. viable Furthermore the results from the tensile tests suggest that parts made by HSS should have sufficient mechanical properties for a wide range of applications. It should be stressed that in addition to the HSS process described in this paper there are other processes such as Speedpart (4), Selective Inhibition Sintering (5) and Electrophotographic Layer Manufacturing (6) that employ the concept of processing entire layers of powdered material simultaneously rather than by a scanning laser. All of these processes, and others using similar concepts will be likely to achieve significant reductions in machine cost, build time and machine cost per part when compared with SLS.for series manufacturing volume significantly greaterReferences 1. Hopkinson, N.

gh Speed UV Laser Beam Scanning by Polygon MirrorKenji Yamazawa, Toshiki Niino The Institute of Physical and Chemical Research (RIKEN) Seiji Hayano ASPECT Co. and Takeo Nakagawa University of Tokyo1.IntroductionSince the stereolithograpy system appeared on the market, various rapid prototyping machines based on the layer laminated manufacturing process have emerged one after the other, the range of applications of which is classified according to their features. Of these, stereolithogrraphy methods which make use of photocurable resin as material and laser beam as processing tools are gradually being established. They are currently the most widespread with a rich variety of models available. The scanning device used in stereolithography serves as the key technology for this method but consists basically of only the galvanometer mirror known for outstanding high speed ability and the XY plotter known for outstanding scanning position accuracy1),2).

Lately, many models are seen to use the galvanometer mirror due to improved accuracy, but still this scanning device is used only when the fabrication size allows, etc. From such reasons, the authors have developed a laser stereolithograpy system capable of high speed high accuracy position scanning by the high speed raster scanning method using a polygon mirror as the scanning device3). The polygon mirror, which has a higher scanning speed than the galvanometer mirror, is moved along the NC servo shaft and the raster scanning spacing is made considerably smaller than the laser beam spot diameter. This paper describes this system and the results of fabrication experiments with this system.2.

This paper describes this system and the results of fabrication experiments with this system.2. Outline and Features of SystemThis system is composed of a laser optical device system made up of a laser oscillator, EOM laser modulator as laser shutter synchronizes with X axis and Y axis, beam expander, and reflection mirror, optical scanning device system made up of the polygon mirror, and f elens , recoating device, liquid resin supply unit, resin tank, and system controller. Its main features include a high speed scanning speed by the polygon mirror of 138.2m/s and 69.2m/s, high position accuracy scanning of below + 0.05mm at the resin surface, use of highly viscous liquid resin of approximately 2200cP with a unique recoating method, and highly accurate fabrication with a layer thickness of 0.05 to 0.2 rom.

The system has a simple and stable mechanism; the resin is drawn from the resin tank by the bucket, poured over the fabricated part and the liquid level is controlled by an overflow method. The 3D CAD data slice data generation software used is SOUPware by NTT DATA CMET INC. The slice data edited on the software is transferred to the controller of the unit from EWS by Ethernet and converted to raster scanning data. The photocurable resins used are mainly the ASAHI DENKA KOGYO 's epoxy resin HS series and versions ofthese resins with high viscosity. Table 1 shows the specifications of the system, Fig. 1 shows the configuration of the system,223and Fig.2 shows the external view ofthe system and controller. a e.

Configuration of the stereolithography system using by polygon mirrorFigure 2. External view of the system and controller2243. Micro Spacing High Speed Raster ScanningThe polygon mirror is a laser beam deflection device currently used in scanning optical system such as the laser printer, etc. As shown in Fig.3 , a polyhedron mirror is rotated at a constant speed by a high speed motor to reflect the laser beam by each of its face so that scanning is performed linearly in one direction in one dimension4). Consequently, it is incapable of vector scanning which traces the contour of the slice data along one line during fabrication. Since the polygon mirror is rotated at an equiangular speed, the angular changes of the reflected laser beam are proportionate to time. A f8 lens is used to scan at a constant speed. For the polygon mirror to rotate stability, it cannot be rotated at low rotational speeds of less than about 1000 rpm.

A f8 lens is used to scan at a constant speed. For the polygon mirror to rotate stability, it cannot be rotated at low rotational speeds of less than about 1000 rpm. Scanning can therefore be performed only at high speeds, and the resin does not reach the critical exposure required for it to harden with only one scanning line at the resin surface. From such reasons, the scanning optical system incorporating the polygon mirror, etc. is moved by NC servo driving at very small spacing compared to the laser beam linewidth (spot diameter) as shown in FigA while performing high speed raster scanning to form a consistent cured layer by drawing the desired photocurable liquid resin surface and average exposure.Cylindrical lensRefleetion mirrorL8serbeamFigure 3. Optical scanning systemLoser beam linewidth (100 j.Lm)OffsetCured linewidth (200 j.Lm)OffsetRoster scanning line pitch <0.33 j.L m 20 j.L m)II01 C 'c c o u (/)c o 01 >, o 0->Lx (Recoot)Figure 4.

Drawing of micro spacing raster scanning locus2254. Recoating SystemThe scanning optical system comprised of the polygon mirror and the recoater which levels the unhardened resin liquid layer have been integrated and are moved together along the same axis to realize recoating of the liquid layer and raster scanning simultaneously. The liquid surface is controlled in a different way in this system. As shown in Fig.5 ,"the semi controlled liquid surface method" is used, in which laser beam is irradiated over the recoated layer, whose liquid surface has been semi controlled, immediately after the recoater blade passes through. This eliminates the need for waiting after recoating in order to remove the effects of the resin viscosity and trapped volumes of the fabricated product on the liquid level.Direction moved by recoater < Recoater bladePlane surfaceLiquid resinSolidified partFigure 5. Semi controlled liquid surface method5.

Semi controlled liquid surface method5. Relation Between Laser Beam Raster Scanning Spacing (Raster Scanning Line Pitch), Recoating Speed, and Cure DepthThe cure depth required for the desired layer thickness can be calculated with the following equations from the laser beam scanning spacing, recoating speed, average exposure, and features ofthe resin used. The relation of these factors is shown in Fig. 6.Average exposureEav [mJI cm2] = PI. I (Lp • Vs • 10) 5)where; PL : Laser output on liquid resin surface[mW]Vs : Laser scanning speed [m/s] (set according to the rotation speed of the polygon mirror) Lp : Raster scanning line pitch [mm](set according to the recoater speed Vr)Cure depth Cd [mm] = Dp.

In ( Eav I Ec) 5)where; The critical exposure is Ec = 7.9[mJI cm2] and penetration depth is Dp = O.08[mm] based on "the Working curve" ofthe photocurable resin.In this system, the raster scanning line pitch is made large and the recoating speed is raised as shown in Fig.7 in order to shorten the time required to laminate one layer. This however results in a shallow cure depth. Consequently, to shorten the fabrication time, it is necessary to project the average exposure over the liquid resin surface in such a way that the required cure depth can be obtained in respect to the layer thickness set, which means that the required cured depth can be achieved by using a high laser output and raising the photosensitivity.6. Fabricating ProcessAs shown in Fig .8 , this unit completes the lamination of one layer by drawing the resin with the bucket, pouring it over the fabricated part, and performing recoating and laser scanning226simultaneously.

As seen in Fig. 11, the surface of the part fabricated by the micro spacing high speed raster scanning method and properties of the contour are good, indicating no need for contour scanning such as vector scanning. The minimum cured linewidth can be set on the control data up to 0.02 mm. As shown in Fig. 12, the linewidth could be calculated to the minimum ofO.lmm. This linewidth is the same as or smaller than that obtained by vector scanning using a galvanometer mirror. The cross-section of the part shown in Fig. 13 indicate that the part can be fabricated according to the thickness set. Lamination was also found to be good with the high viscosity resin (approx. 2200cP) and no effects of the change in the resin viscosity on the fabrication process were seen in particular. When the thickness of the uncured liquid is greater than the layer thickness set (cure depth), peeling occurs between the layers during fabrication.

When the thickness of the uncured liquid is greater than the layer thickness set (cure depth), peeling occurs between the layers during fabrication. This requires the position of the recoater blade bottom to be set at an accuracy of +O.lmm in respect to the liquid resin level.. xampJeo amma mg con lIon .Iffl Layer thickness Viscosity of resin Beam scanning velocity Raster scanning line pitch Recoating velocity Average exposure Critical exposure Penetration depth Cure depthT: bl 2 Ea eO.lmm 80-2200 cP 138.2m/s 0.83 j.t m 0.28mm/s 56.44 mJ/cm2 7.9 mJ/ cm2 0.08mm 0.16mmFigure 9. Photograph of the fabricated sample partEA~IFroI========t:=r======::::J~~Table 3.

Photograph of the fabricated sample partEA~IFroI========t:=r======::::J~~Table 3. Accuracy (Average of 10 samples: rmml)Position Dimension MeasurementA B C D E F G H I J K L 21 22 23150.00 150.00 50.00 60.00 15.00 15.00 15.00 15.00 15.00 15.00 15.00 15.00 25.50 15.00 5.00150.01 149.96 49.96 60.00 15.01 15.04 15.05 15.05 15.05 15.03 15.03 15.05 25.50 15.08 5.04Error +0.01 -0.04 -0.04 0 +0.01 +0.04 +0.05 +0.05 +0.05 +0.03 +0.03 +0.05 0 +0.08 +0.04Figure 10. Drawing of sample part228yLx O.lmmLJFigure 11. Photograph of surface and outline ofpart (Viscosity of resin 2200cP)yLx O.lmmLJFigure 12 . Photograph of minimum cured linewidth ofpart (Viscosity of resin 2200cP)z Lx O.lmmLJFigure 13. Photograph of cross section of part (Viscosity of resin 2200cP)2298. Summary and Future TasksA new stereolithograpy system which employs a polygon mirror laser scanner and performs recoating and scanning simultaneously has been developed.

Summary and Future TasksA new stereolithograpy system which employs a polygon mirror laser scanner and performs recoating and scanning simultaneously has been developed. The fabrication results show that layer manufacturing can be carried out for a broad range of resin viscosities using this high speed high accuracy scanning method. The fabrication process is also useful for reducing the waiting time after recoating in stereolithography process. In the future, efforts will be made to quantify the optimum fabrication parameters such a layer thickness, coefficient of resin viscosity, recoating speed, laser irradiation position, and cure shrinkage to raise the accuracy of the part built and put the method to practical application. Efforts will also be made to realize more accurate lamination of below 50 j..J. m.

Efforts will also be made to realize more accurate lamination of below 50 j..J. m. Finally, the use of such laser beam deflecting devices capable of high speed scanning like the polygon mirror for stereolithography may shorten the fabrication time remarkably, given that lasers with higher outputs than current laser are available or photosensitivity of the photocurable resin is raised. Considered a promising stereolithography method, efforts will thus be made to further optimize the system as a productive and practical unit.AcknowledgementsThe authors would like to thank NTT DATA CMET INC. the co-developers of this device, aswell as all those who gave their cooperation.References1) Y. Marutani, K. Ookawa, S. Hayano, N. Saito, and T. Nakai, "Hikari-zokei-hou (Laser Stereolithography - 3D-plotter using laser)", Nikkan Kogyo Shinbunsha, Japan, 1990, ppl08 109 (in Japanese).2) T. Nakagawa, Y.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperHIGH SPEED VIDEO IMAGING OF OVERHANG SURFACES IN BEAM SHAPED LASER POWDER BED FUSION OF 316L STAINLESS STEELL. Vanmunstera,b, A. Dejansa, B. Van Hoorewedera, B. Vranckena,ba) KU Leuven, Department of Mechanical Engineering, Celestijnenlaan 300, 3001 Leuven b) FlandersMake@KU LeuvenAbstract Beam shaping has seen rising interest in the AM industry and research field because of the increased level of control over the spatial distribution of the thermal input during Laser Powder Bed Fusion, allowing for faster build rates. This report investigates if the more uniform heat input offers benefits to the scanning of unsupported overhanging structures. In-situ high speed video imaging is combined with post process surface characterization to derive a relation between the scan parameters and quality of the overhang surface.

In-situ high speed video imaging is combined with post process surface characterization to derive a relation between the scan parameters and quality of the overhang surface. The best results were achieved at low energy densities, but no clear advantage of top-hat shaped laser spots over the standard Gaussian beams is observed.Keywords: High Speed Video, Beam shaping, Overhang, Laser Powder Bed Fusion****************************************************************************************** IntroductionLaser Powder Bed Fusion (LPBF) is a metal Additive Manufacturing process capable of producing highly intricate parts by translating a computer 3D model into layers, which are consequently selectively molten by a high power laser on a fine metal powder bed. By designing the optics train to produce small laser spot sizes, small melt pools and hence small minimal feature sizes are obtainable.

By designing the optics train to produce small laser spot sizes, small melt pools and hence small minimal feature sizes are obtainable. However, the small laser spot and thin metal layer thickness, required for good part accuracy and surface quality, severely limit the maximum process speed. One of the mitigation strategies under investigation for this issue is the use of larger spot sizes combined with higher laser powers. However, a mere defocusing of a standard single-mode (Gaussian) laser spot leads to a quickly deterioration beam quality and hence less reliable part qualities [1]. Therefore, different beam profiles with a more uniform heat input are under investigation. These are amongst others Top-hat shaped, Bessel beam shaped or Ring shaped, allowing to achieve consistent temperature distributions across a wider area and reduced temperature gradients [2].Literature suggests the ring-shaped spots lead to shallower, wider and more stable melt pools [3].

The increased width could lead to a larger process window and hence a significant productivity increase through the use of higher scan speeds and larger hatch distances [4], [5]. However, the larger spot sizes inherently expand the minimal achievable feature sizes. Hence, OEMs try to include multiple beam sizes in a single system, e.g. through the use of multiple scanners or with variable focusing units. Another option is to install a laser with integrated beam shape switching, such as the AFX-1000 laser by nLight inc. [6] which allows switching between a small single-mode shape and a larger multimode ring-shape in between individual scan tracks.While the different beam shapes are mostly under investigation for their potential gains in productivity, they can also be beneficial to other part qualities. The more uniform energy input might allow for enhanced properties of the overhanging and downfacing areas that are currently a major challenge for conventional beam shapes.

The more uniform energy input might allow for enhanced properties of the overhanging and downfacing areas that are currently a major challenge for conventional beam shapes. In these regions, the laser scans over areas of loose powder without any bulk metal support. The highly reduced thermal conductivity of powder compared to dense metal severely limits the ability of the material to spread the input energy. This results in larger melt pools and increased temperatures, which in turn lead to higher residual stresses and the formation of dross [7]. The latter is a sinking of the melt pool of overhang layers in the powder bed under the influence of capillary forces and gravity and results in a loss of dimensional accuracy [8]. Secondly, the layers exhibit a high level of internal stresses, which can lead to upwards warping and consequent recoating issues.

Secondly, the layers exhibit a high level of internal stresses, which can lead to upwards warping and consequent recoating issues. Typically, this is accounted for by adding support structures underneath overhanging areas, to act as a mechanical anchor and heat dissipator. However, the scanning of support structures increases the build time, leads to material waste and increases the postprocessing time, since they have to be removed after production1105[9]. Another strategy was investigated by Kruth et al., who implemented feedback control to achieve reduced dross levels [10]. Ashby et al. [11] combined high speed optical imaging with infrared (IR) imaging and simulations to better understand the effect of varying scan parameters on the overhang quality.

Ashby et al. [11] combined high speed optical imaging with infrared (IR) imaging and simulations to better understand the effect of varying scan parameters on the overhang quality. Their work concluded that a strategy for reduced energy input in the overhanging area is required to prevent material aggregation due the overheating.This work investigates the potential of a beam shaped laser to improve the dimensional accuracy of overhanging areas in LPBF parts. Post-process surface quality analysis will be combined with in-situ high speed optical imaging to study the physics behind powder agglomeration, dross formation and layer warping.Methods & MaterialsSample DesignThe overhanging areas were studied using bridge-like structures supported on two sides. The overhang samples consist of a 0.5 mm thick bridge build on top of two substrate pillars. All dimensions are displayed in Figure 1.Figure 1: Design and dimensions of the bridge samples.

The overhang samples consist of a 0.5 mm thick bridge build on top of two substrate pillars. All dimensions are displayed in Figure 1.Figure 1: Design and dimensions of the bridge samples. The build direction BD is oriented along the Z axis, while X is the coater direction.Laser Powder Bed FusionAll samples were produced on a heavily modified Concept Laser M1 machine that was retrofitted with a AFX-1000 Variable Beam Profile laser from nLight inc. This laser allows for fast switching between multiple beam profiles by dividing the laser power between the single-mode core and multimode-enabling fiber cladding, resulting in a combination of a Gaussian and a ring profile. Seven discrete levels of power distribution between the Gaussian peak and surrounding ring are possible, ranging from index 0 (near 100% single-mode or Gaussian beam) to index 6 (90% of power in the ring). It has a maximum power rating of 600 W in the Gaussian peak and 1200 W in the ring and a 1070 nm wavelength.

It has a maximum power rating of 600 W in the Gaussian peak and 1200 W in the ring and a 1070 nm wavelength. To accommodate beam switching, the laser and scanner control was transferred to an in-house developed control system, allowing for in-layer beam shape switching. A large aperture Scanlabs scanner and a f-theta lens with focal distance of 420 mm were implemented in the optics train. Due to a defocusing issue in the setup, the Ring shape was not achieved, but rather a 250 µm top-hat shape for index 6 was measured, while the Gaussian-shaped index 0 spot was measured to be 150 µm wide. All samples were built using 316L stainless steel powder from Höganäs (Sweden) with a 15 to 45 µm particle size range, under an Argon atmosphere (gas flow in the Y direction of Figure 1). To reduce the likelihood of a warped sample damaging the recoater, all samples were oriented with their long axis in the recoater direction (X axis in Figure 1).

To reduce the likelihood of a warped sample damaging the recoater, all samples were oriented with their long axis in the recoater direction (X axis in Figure 1). A zigzag (meander) scan strategy without contours was used, where the hatching orientation is altered by 90° for every subsequent layer. By using a fixed layer thickness set to 30 µm, a total of 17 overhanging layers were scanned. The first overhang scan is oriented in the X direction (see Figure 1), perpendicular to the substrate pillar edge as this is thought to be preferential for part quality [10].In-situ high speed optical imagingFor high speed optical observations, a Phantom VEO 640 high speed video camera was used, combined with an Infinity Model K2 DistaMax long range microscope lens. Videos were captured at a frame rate of 18000 Hz with a 53.8 µs exposure time and a resolution of 512x384 pixel. During recording, an LED lighting system of approximately 50000 lux was focused on the recorded area.

During recording, an LED lighting system of approximately 50000 lux was focused on the recorded area. Although the camera and lens were positioned outside1106of the Concept Laser M1 machine, a mirror set up in the build chamber ensured a viewing angle perpendicular to the build plate. The overall setup resulted in a field of view of approximately 5.2x3.9 mm2, equivalent to a pixel size of just over 10x10 µm2 .Post-situ surface characterizationFor imaging of the overhang surface after production, as well as for the dross depth measurements, a Keyence VHX-6000 Digital Microscope was used. By applying a 100x magnification, a 3D image of a 3.5x2.5 mm2 surface in the center of the overhang area is measured using a Depth of Focus method.

By applying a 100x magnification, a 3D image of a 3.5x2.5 mm2 surface in the center of the overhang area is measured using a Depth of Focus method. An in-house developed Matlab code was used for the image processing, such as creating an average profile along the scan direction.Results & DiscussionDimensional accuracy of overhang regionTo examine the effect of varying the laser power P, scan speed v and hatch spacing h, as well as the spot shape and size, 100 bridges were built with varying parameter combinations. Figure 2 gives an overview of the parameter ranges examined. Only the index 0 and index 6 beam shapes were used, and will be referred to as ‘Gauss’ and ‘Top-hat’, respectively. All samples used equal parameters for the substrate pillars (P = 285 W – v = 800 mm/s – h = 95 µm – Gaussian shape), only varying parameters in the 17 overhanging layers.Figure 2: Overview of the investigated parameter ranges.

Every combination of P, v & h was built once with index 0 (Gauss shape) and once index 6 (top-hat shape).For every bridge, the resulting height of the overhanging area was measured by averaging 30 measurements of the height difference between the center of the bridge’s downfacing surface and the top surface, as schematically presented in Figure 3a. The resulting height measurements are shown in Figure 3b relative to the Volumetric Energy Density (VED) of the overhang parameters, which is calculated as: 𝑉𝐸𝐷 =𝑃 ℎ 1000𝐽𝑚𝑚3] [(1)𝑡 ∗ 1000 With P the laser power [W], v the scan speed [mm/s], h the hatch spacing [µm] and t the layer thickness [µm].𝑣 ∗Figure 3: Measurement of bridge height: a) example of the measurement location; b) resulting measured bridge heights for samples with small hatch spacing, where the design thickness was 500 µm; c) resulting measured heights for sample with large hatch spacing, where the design thickness was 500 µm.

Samples that did not have a flat top surface were not included in the figure.1107Not all parameters lead to defect free parts. Generally, two types of failure are observed. Firstly, when the input energy density is lower than 50 J/mm3, insufficient melting occurs to achieve a solid top layer. Secondly, most failures occur due to a full or partial separation in the bridging surface layer. This warping effect is generally attributed to internal stresses in the layer as a result of rapid and uneven heating/cooling cycles. This upwards curling of the surface prevents proper application of the following powder layer, which further hinders the development of a solid top layer.The measurement results are summarized in Figure 3, excluding the failed samples. Across all samples, a clear increasing trend between the energy input and bridge height is observed.

Across all samples, a clear increasing trend between the energy input and bridge height is observed. Since each bridge is composed of the same amount of layers, the variation in bridge height is not related to the CAD design, but to other effects: capillary forces cause dross formation and gravity causes a sinking of the molten material in the powder bed. A decrease in input energy density leads to a decrease in melt pool size, reducing the time the material is in a liquid state, thus hindering capillary workings and hence a decrease in bridge height is observed [7]. This explains the relation between the observed bridge height and energy input. When comparing the effect of the laser shape, two observations are made:In the parameter combinations where the hatch spacing was smaller than approximately half the top-hat laser spot size (i.e. <120 µm), the Gaussian index 0 beam leads to thinner bridges, even at equivalent energy levels.

<120 µm), the Gaussian index 0 beam leads to thinner bridges, even at equivalent energy levels. It appears that, even though the index 6 top-hat shape spreads the energy over a larger area, this increases the time the material is in molten state. Hence the melt beads could grow larger and penetrate deeper in the powder bed. This could be the result of the remelting that occurs in every consecutive melt track, since there is considerable overlap between the tracks of consecutive scan vectors. Figure 4 is an illustration of this remelting, where the orange remelting zone overlaps with the previous scan track. • When comparing the parameters with larger hatch spacing (i.e. ≥120 µm), there is no discernable difference in behavior between the index 0 and index 6 beams. Overall, the spread on the results is larger in the case of larger hatch spacing, potentially owing to an increased denudation effect.

Overall, the spread on the results is larger in the case of larger hatch spacing, potentially owing to an increased denudation effect. Still, the resulting bridge heights at equal energy densities are lower in the large hatch cases. Presumably the remelting in every step discussed above has a stronger effect on the bridge size than the level of VED.Figure 4: Schematic representation of the remelting due to small hatch spacing and large spot size: a) small Gaussian laser spot with small hatch spacing. The orange remelting zone does not overlap with the previous scan track; b) large top-hat shaped spot with small hatch spacing, exceeding the distance between consecutive tracks. A significant remelting zone is indicated in orange, with clear remelting of the previous track.High Speed Video ImagingFor some parameter sets, the processing of the first layer of the overhang was recorded using a stationary off-axis high speed imaging system.

An example of an image acquired in this way is shown in Figure 5. The goal of these videos is to help visualizing the difference in process behavior in the transition from dense material (i.e. the substrate pillar) to scanning on loose powder (the overhang region). The lower thermal conductivity greatly reduces the process stability, with the videos showing a visible increase in the amount of spatter and denudation. The melt pool is no longer continuous, but rather forms non-connected or loosely connected individual droplets, which is known as balling. This effect is the result of surface tension of the liquid melt pool that can no longer be counteracted by adhesion to a previously solidified layer [12]. This was also partially observed by Ashby et al. [11], albeit in this work at much lower power levels than they used, leading to smaller bead sizes.

This was also partially observed by Ashby et al. [11], albeit in this work at much lower power levels than they used, leading to smaller bead sizes. The overhanging1108part of the sample also glows for several hundreds of ms after the passage of the laser, which is again a direct result of the low thermal conductivity of the powder. This effect was measured as the time it took a specific region to reach stable intensity values. While the exact glow time depends on lighting conditions, camera properties and settings, a relative comparison between samples can be made. On average a 192 ± 33 ms afterglow was measured on the high speed videos in the center overhang part, while no visible afterglow is noticeable in the solid substrate. Hence, suggesting the cooling rate in the first overhang layer is orders of magnitude lower compared to the substrate pillar. The afterglow measurements are shown in Figure 6.

Hence, suggesting the cooling rate in the first overhang layer is orders of magnitude lower compared to the substrate pillar. The afterglow measurements are shown in Figure 6. The afterglow time is found to correlate well with the VED, but did not show a statistically significant difference between Gaussian laser and top-hat laser samples at equal scan parameters. The effect of thermal conductivity can also be seen by comparing afterglow measurements in the overhang center (indicated as point C in Figure 5c) to those in the transition between substrate and overhang edge (point E). Point C is located about 2 mm from point E. On average, the transition region E cools down about 45% faster than the overhang center C, likely showing the effect of the substrate pillar as a heat sink.

On average, the transition region E cools down about 45% faster than the overhang center C, likely showing the effect of the substrate pillar as a heat sink. Lastly, the scanned layer visibly sinks in the powder bed during cooldown, thereby aggravating the dimensional deviation of the sample compared to the CAD design.Figure 5: Still frame from the high speed imaging showing the difference in melt quality atop the substrate pillars and overhanging powder layer: a) schematic showing the position of the image in red; b) during the scanning process. Some glowing metal beads are indicated in red, spanning more than 1 scan track in size; c) after scanning and cooldown.

Some glowing metal beads are indicated in red, spanning more than 1 scan track in size; c) after scanning and cooldown. Locations of afterglow measurements are indicated by the orange arrows for the overhang center (C) and overhang edge (E).Figure 6: Measurement of visible afterglow due to elevated temperature and low powder thermal conductivity for locations measured on the overhang edge or in the overhang center: a) afterglow measurements per parameter combination. The location of edge and center are indicated in figure 5c by points E and C respectively; b) afterglow compared to energy input.Figure 7 shows a comparison between the first layer behavior of the index 0 and index 6 laser spots for three different VED levels. An increase in energy input leads to increased bead sizes, as is also observable in the boxplots of the bead surface areas on the right side of the figure. This behavior is true for both laser spot shapes under investigation.

This behavior is true for both laser spot shapes under investigation. When comparing index 0 and index 6 first layers at equivalent energy input levels, an increase in average bead surface area is also noticed. This increase is dependent on the energy input level and ranges from +30% to +60% larger bead size when switching from a smaller Gaussian spot to a larger top-hat, while keeping the other scan parameters constant. Hence, the larger spot size presumably allows the material to stay in molten state for a longer time, increasing the probability of larger bead formation. In Figure 7f, the powder bed is still1109glowing at the end of the recorded video. Even though the energy levels per mm2 are similar between figures (e) and (f), the slower scan speeds increase the laser-material interaction time and hence might result in much larger beads.

Even though the energy levels per mm2 are similar between figures (e) and (f), the slower scan speeds increase the laser-material interaction time and hence might result in much larger beads. As these image were taken for parameter sets with smaller hatch spacing, the increased bead size could be a cause for the difference in bridge thickness when varying the laser spot shape observed in the previous section.Figure 7: Overview of overhang layers after the first scan, with varying parameter sets. Figure f was still glowing at the end of the recording due to excessive energy input and slow scan speeds. Box plots shows the bead size measurements for each parameter set a – f, with X representing the average bead size.Interestingly, a certain periodicity develops in the beads on loose powder for some parameter combinations.

Subsequent tracks fuse preferentially to the beads formed by the previous scan track, which causes a growth of the beads in the direction perpendicular to the scan vectors. Figure 8 illustrates this behavior. This could result from a combination of the beads acting as a heat sink and adhesion between currently molten and previously solidified material. Since the coalescence of molten beads draws material particles from the surrounding powder bed, the areas in between beads also form a line perpendicular to the scan direction. An example is highlighted in Figure 9a. The effect most likely leads to an increased risk of layer warping. These line- like powder free zones between the beads could act as mechanically weak zones in the layer due to their lack of material. The risk of cracking might be aggravated even further because the weak lines are oriented approximately perpendicular to the direction of largest interior cooling-related stresses [13].

The risk of cracking might be aggravated even further because the weak lines are oriented approximately perpendicular to the direction of largest interior cooling-related stresses [13]. From the videos, it can be observed that the layer breaks along these weak zones during cooling, resulting in individual “islands” of fused material surrounded by cracks. These islands are seen the move independently of each other during cooling. Figure 9 is used to indicate the size of the individual islands that were observed.For the parameter combination shown in Figure 9, a remelting step was performed approximately five seconds after the first melting scan, to try to fuse the individual solidified islands together. This was only tried for one parameter combination, as an initial estimate for the usefulness of remelting for overhangs.

This was only tried for one parameter combination, as an initial estimate for the usefulness of remelting for overhangs. During cooldown, the formation of at least one crack was still visible, as shown in (c), but the amount of cracks is reduced, since the individual islands are fusing together. Conversely, this remelting step also visibly increased the melt bead size and amount of layer sinking in the powder bed. Hence, no further investigation of remelting was performed.1110Figure 8: Crack growth due to coalescences of consecutive molten beads. Between every image one extra vector is scanned. The black lines are material lacking zones in between de molten beads, clearly showing growth approximately perpendicular to the scan tracks.

Between every image one extra vector is scanned. The black lines are material lacking zones in between de molten beads, clearly showing growth approximately perpendicular to the scan tracks. Hence, the molten beads also form horizontal lines.Figure 9: Still frames of parameter combination P = 150W, v = 800mm/s, h = 95µm, index 0: a) right after melting, the overhang surface still glows and shows the melt beads coalescing, leading to line like weak zones; b) after cooling, 2 cracks formed in the solidified overhang layer. The resulting islands are seen moving independently of each other; c) after remelting and cooling one crack is still present.Post-process Surface Quality AnalysisTo validate the findings from the video analysis, the downfacing surface of the successfully printed parts were optically scanned using a Depth form Defocus method to form a 3D representation of the surface, an example result of which is shown in Figure 10.

From these images a height map can be extracted to study the dross beads and intermediate weak zones discussed in previous paragraphs. The level of connectivity of the dross beads could be a good measure for the thermal and mechanical support a layer provides for the next one. The degree of connectivity can be studied by calculating the average profile of the measured surface perpendicular to the scan direction. This is done by summing all data points in the Y direction and normalizing the result, an example of which is shown in Figure 10c. Any continuous material lacking zones show up as sharp valleys in the averaged profile.1111Figure 10: Analysis of the resulting downfacing surface of parameter set [150W, 1000mm/s, 85µm, single-mode]: a) 3D optical image at 100x zoom; b) height map of the 3D image in (a); c) averaged profile along the Y direction.

Red lines are used to indicate continuous deep valleys in the 3D surface and averaged profile.Figure 11 show the height maps and averaged profiles for three parameter sets. The medium-low VED level in (a) and (d) shows a low degree of periodicity, leading to a successfully printed sample. Meanwhile, sharp peaks in the averaged profile of the low VED sample in (b) and (e) indicates the presence of multiple weakened zones in the top surface, with valley more than twice the layer thickness used. This sample failed due the layer warping and insufficient melting. The sample in (c) and (f) does not show deep weak zones, but the high energy input in this high VED sample lead to a significant curvature of the bridge’s bottom surface towards the build plate, which indicates sinking of the material in the powder bed.In order for a layer to fail due to warping, two conditions need to be satisfied.

Firstly, there has to be a significant stress level in the sample and secondly the underlying material layer must be too weak to carry this stress. Thus, a bridge should be sufficiently thick and have as few line-like cracks in the surface as possible. Four quality metrics were be examined for a good parameter combination: (1) it should lead to good dimensional accuracy of the bridge height, (2) to a low degree of curvature in the bridge bottom surface, (3) the absence of weaker zones that could cause warping under stress buildup, and finally (4) a fast recuperation to bulk processing quality. Figure 12 gives an overview of these 4 metrics for all samples. Here, the samples are divided in 3 groups based on the quality of the top surface, i.e. how well a parameter set can recuperate to acting as a solid bridge after 0.5 mm of overhanging layers. Good quality parameter combinations lead to flat top surfaces without any remaining signs of the overhanging area.

Good quality parameter combinations lead to flat top surfaces without any remaining signs of the overhanging area. Sufficient quality samples are almost flat, showing a slight curvature in the bridge outer edges, but are most likely recuperating to a fully flat surface if the build job would be continued. Bad quality samples either did not lead to a bridge in one piece or have such a bad quality top surface that further building is not expected to not lead to dense samples.Figure 12a shows the bridge height compared to VED, as already discussed earlier, but also includes the top surface quality grouping. Below 50 J/mm3, no good quality parameter combinations were found, indicating an insufficient energy level for good fusing of material.

Below 50 J/mm3, no good quality parameter combinations were found, indicating an insufficient energy level for good fusing of material. Conversely, the test did not show good quality bridges above 118 J/mm3, with failures probably linked to high thermally induced stresses as a result of excessive heat input.In Figure 12b, the deepest material lacking zone for every average profile is plotted. There is no clear link between the energy input level and the depth of these zone. There is however a good correlation between the level of depth and the chance a sample fails to achieve a good quality top surface. No good quality samples were built if a crack deeper than 60 µm (i.e. twice the layer thickness) is present. We can thus use this metric as a measure for how close the sample was to failing due to warping.

Here, a negative curvature means the bottom surface of the bridge curves downwards, i.e. towards the build plate. Samples with a high energy input curved downwards, which is linked to a high degree of dross formation, but also to the sinking of the powder bed as shown in the high speed video imaging and was also observed in Figure 11f. The high energy levels and corresponding long interaction time between liquid metal and powder particles directly underneath exaggerate the sinking. This effect opposes the expected upwards curling due to thermal stress. A decrease in energy input induces reduced levels of sinking and hence reduced curvature. At the lowest levels of VED tested, the average curve is directed upwards, indicating that warping forces are greater than the sinking effect.When an optimized overhanging parameter set needs to be defined, the information in the graphs described above can be combined.

The best quality parameters lead to low bridge heights (VED slightly above 50 J/mm3 in our experiments), low bottom surface profile depths and low curvature values. For the Gaussian beam shape samples, this lead to one best set: P = 150 W, v = 1000 mm/s & h = 95 µm, which is a VED level of 53 J/mm3. For the top-hat beam shape samples, two different sets are identified. When optimizing for bridge height this was P = 285 W, v = 1000 mm/s & h = 120 µm (79 J/mm3), while the lowest curvature was achieved at P = 150 W, v = 1000 mm/s, h = 85 µm (59 J/mm3). Note that the latter set for index 6 is almost the same parameter set as the optimal index 0 setting. This confirms that, at least in the parameter ranges and optical setup investigated in this report, no noticeable advantage of larger top-hat laser beams compared to small Gaussian beams was found.

This confirms that, at least in the parameter ranges and optical setup investigated in this report, no noticeable advantage of larger top-hat laser beams compared to small Gaussian beams was found. However, if larger parameter ranges, especially larger hatch spacing are examined, a difference might still arise.1113Figure 12: Bridge dimensional quality metrics versus VED: a) Bridge height versus energy density, cfr. in figure 3. Useful energy levels for overhang scans range from 50 – 120 J/mm2; b) Maximum valley depth versus VED. With one exception, all good parts have a valley depth below 60 µm; c) downfacing surface curvature versus VED, showing an inverse relation.ConclusionIn this report in-situ high speed imaging and post-process surface characterization are used to study the Laser Powder Bed Fusion (LPBF) process in an unsupported overhang situation.

The focus lays especially on comparing the effect of a change in spot shape and size on the bridge quality:The dimensional accuracy of overhang bridges was measured by characterizing the levels of dross and material sinking with varying laser power, scan speed and hatch spacing. A clear relation between the energy input and depth of material sinking is observed.High speed videos showed how the discontinuous melt pool coalesces in the beads perpendicular to the scan track.

A clear relation between the energy input and depth of material sinking is observed.High speed videos showed how the discontinuous melt pool coalesces in the beads perpendicular to the scan track. The first scanned layer was found to not necessarily be a solid layer, but rather a series of individual fused islands surrounded by cracks.Optical imaging of the resulting overhang surfaces were used to study the quality metrics of a sufficient overhang part.The optimal parameter sets for dimensionally accurate overhangs for Gaussian beam and top-hat beam lasers are almost identical in the investigated parameter range, indicating no immediate benefit of the top- hat shape when the hatch spacing values are similar or lower than its spot size.

merical simulation and experimental investigation of arc based additive manufacturing assisted with external longitudinal static magnetic fieldX.M. Zhou*, H.O. Zhang*†, G.L. Wang** , L.Y. Liang**, Y.H. Fu**, X W Bai* and X.P Wang*State Key Laboratory of Digital Manufacturing Equipment and Technology, Huazhong University of Science and Technology, Wuhan 430074, PR China **State Key Laboratory of Materials Processing and Die & Mould Technology, Huazhong University of Science and Technology, Wuhan 430074, PR ChinaAbstractThis paper proposes a new arc based additive manufacturing method assisted with external longitudinal static magnetic field. An electromagnetic coupling numerical model has been established, which consists of droplet impingement, heat transfer and dynamics of molten pool.

An electromagnetic coupling numerical model has been established, which consists of droplet impingement, heat transfer and dynamics of molten pool. Comparing the simulation results between normal deposition and external longitudinal magnetic filed assisted deposition, it shows that the external longitudinal static magnetic field induces the tangential stirring force in molten pool and drives the molten metal moving to the edge of the pool. This can reduce the temperature gradient in forming region. Furthermore, the related comparison experiments of single-bead deposition and multi-beads overlapping deposition are performed, it was found that the tangential stirring force can reduce the height of single-bead as well as increase the width of single-bead. The experimental results are in accordance with the simulation results.

The experimental results are in accordance with the simulation results. The changing of morphology of the bead is beneficial to multi-beads overlapping and capable of improving the surface accuracy for arc based additive manufacturing parts.IntroductionArc based additive manufacturing(AAM) is a wire and arc welding based additive manufacturing(AM) technology which has high energy efficiency [1] and deposition rate [2] comparing with laser or electron beam deposition AM technology. In AAM process, components are fabricated by overlapping weld beads in one layer and deposition of materials layer-by-layer in height direction，however, it is easy to introduce geometric error due to the accuracy limitation of Gas Metal Arc Welding(GMAW)(or Gas Tungsten Arc Welding, GTAW) and cumulative error in the process of overlapping. Several researches have been devoted to improve surface accuracy for arc welding based AM.

Several researches have been devoted to improve surface accuracy for arc welding based AM. Xiong et al.[3] established a model of bead section profile and overlapping beads with experimental validation for robotic GMAW-based AM, and obtained smooth surface with no defect by determined optimum model for single bead section. Later, Xiong et al.[4] developed a passive vision sensor system to monitor the nozzle to the top surface distance(NTSD) in arc welding based AM, and designed an adaptive control system to keep the NTSD constant by moving the working flat and adjusting the deposition rate on next deposition layer. These strategies can improve process stability and obtain smooth thin-walled parts. Ding et al.[5] built a single weld bead model through various curve fitting methods. Then they established a tangent overlapping model (TOM) and presented a concept of critical centre distance for stable multi-bead overlapping processes.

Then they established a tangent overlapping model (TOM) and presented a concept of critical centre distance for stable multi-bead overlapping processes. The multi-bead and multi-layer deposition experiment indicated that the proposed TOM is better in geometry accuracy and material efficiency than traditional flat-top overlapping model (FOM).The longitudinal magnetic field has been widely employed for electromagnetic assisted arc welding. Malinowski-Brodnicka et al. [6] investigated the effect of longitudinal electromagnetic stirring on austenitic stainless steel GTAW by experiments. It was found that an asymmetric weld bead was formed in the case of a constant magnetic field, when an alternating field was applied, a symmetric bead was formed. The increasing of field strength could result in increasing of weld width and decreasing of penetration. Simultaneously, the application of an alternating magnetic field could lead to grain refining on macro scale as well as micro scale.

Simultaneously, the application of an alternating magnetic field could lead to grain refining on macro scale as well as micro scale. The changing of weld bead shape and metal structure were interpreted in terms of arc expansion, dendrite fragmentation, mixing of the liquid weld metal in the vicinity of the solidification. Chang et al. [7], Jiang et al.156[8] and Luo et al. [9] conducted an experiment of GMAW assisted with an external longitudinal magnetic field and found that the longitudinal magnetic field could improve the characteristics of arc plasma and metal transfer, such as reduction of spatter in certain welding conditions. Chang et al.

Chang et al. [10] studied the impact of the synchronous longitudinal electromagnetic field including low-frequency(5-25Hz) and high-frequency magnetic fields(500-2000Hz) on short-circuit GMAW by multiple sets of comparative experiments, and found that all these two kinds of magnetic fields could improve the frequency of metal transfer and reduce the spatter.To date, the studies of external magnetic field assisted arc welding or weld-based AM are mainly focusing on experimental study. Although the information obtained through experimental observations is very useful, it is difficult to use experimental methods to reveal underlying mechanisms due to high-temperature and transient during the arc welding. Numerical simulation provides an alternative way to obtain insightful physical phenomena. Ushio and Wu [11] built a three-dimensional modeling of heat and fluid flow in a moving gas metal arc weld pool. CAO et al.

Ushio and Wu [11] built a three-dimensional modeling of heat and fluid flow in a moving gas metal arc weld pool. CAO et al. [12] developed a 3-D transient numerical model of gas metal arc (GMA) weld pools with free surface and droplet impact effects by considering not only the heat transfer and fluid flow, then simulated weld bead geometries were in good agreement with experimental measurements. Hu [13] and Rao [14] developed a uniform model covering arc droplet formation, detachment, transfer, impingement and weld pool dynamics. Later, Hu [15] and Rao [16] built a three-dimensional transient model of droplet impingement and weld pool dynamics, and investigated the influence of metal transport phenomena on the formation of ripples. Chang et al. [17] established a finite element model analysis on temperature field and stress field during submerged-arc welding to study the effects on weld bead formation by using longitudinal magnetic field of low- frequency.

It was shown that a molten pool with a wider weld face and a lower penetration could be obtained and by the control of longitudinal magnetic field of low-frequency. Bai et al. [18] presented an induction heating assisted weld-based additive manufacturing process, and developed a coupling analysis FEM model of induction and arc heat. Then induction preheating and postheating on residual stress state are analyzed in comparative simulations. But the simulation model of Chang [17] and Bai [18] didn’t include the weld pool dynamics. Luo et al.[19] utilized a probe to measure the distribution of arc current density with an additional longitudinal magnetic field and developed a model for the fluid flow in the weld pool for the control process, which was based on experiments using LD10CS aluminum alloy. The influences of the additional longitudinal magnetic field and other factors are discussed. Yin et al.

The influences of the additional longitudinal magnetic field and other factors are discussed. Yin et al. [20] established a 3D numerical model containing the welding arc and the weld pool for gas tungsten arc welding (GTAW) with applied axial magnetic fields. The distribution of temperature, heat flux, current density and pressure on the anode surface becomes double-peaked while the applied axial magnetic fields were applied. The fluid in the centre areas rotates in an opposite direction to that in the outer regions. But this model does not include the free surface of weld pool.In this paper, a new arc based additive manufacturing method assisted with external longitudinal static magnetic field is proposed, and a separation numerical model is developed for metal transfer and weld pool dynamics. The schematic of longitudinal magnetic field assisted AAM is shown in Fig.1.

The schematic of longitudinal magnetic field assisted AAM is shown in Fig.1. The detailed comparative analyses of temperature, velocity, and electromagnetic force are performed, and the simulation results are associated with the morphology of weld bead. Then the single-bead deposition and multi-beads overlapping experiments are performed, and the effect of the external longitudinal static magnetic field on the surface quality and dimensional accuracy of the AAM parts are discussed.157Fig.1. A schematic representation of longitudinal magnetic field assisted AAMMathematical models1.

A schematic representation of longitudinal magnetic field assisted AAMMathematical models1. Governing equationsThe governing equations of heat and mass transfer are as following: Mass continuum equation:   t Momentum conservation equation:   t Energy conservation equation:(1)   (V))  (0 (FF stFFFF bV (p      )VV (  )V)(2)magarcmsdrag t  ( h)  (V h)    (k T)SarcSdrop(3)In the above equations, ρ is density, V is velocity vector, p is pressure, and μ is viscosity, and h isenthalpy. The other items are source item and will be given as following.a) Source item of mass In this paper, the droplet is considered generating at a certain frequency and injecting to molten pool.

The other items are source item and will be given as following.a) Source item of mass In this paper, the droplet is considered generating at a certain frequency and injecting to molten pool. The source item of mass can be written as, l  t where ρl is desity of liquid metal, t is droplet generation time.massS(4)b) Source item of momentumThe item Fb is buoyance which can be calculated by Boussinesq approximation (Eq.(5)), where ρ0 is the(constant) density of the flow, T0 is the reference temperature, and β is the thermal expansion coefficient.bF  (  0)g ( T T g 0 0)(5)In current study, the VOF method is employed to track the free surface. F(x,y,z,t) represents the volumeof metallic phase per unit volume of fluid. The VOF equation is expressed as Eq.(444).

F(x,y,z,t) represents the volumeof metallic phase per unit volume of fluid. The VOF equation is expressed as Eq.(444) . When F is between 0 and 1, it means that the cell is located at the free surface, and gradient direction of F that is normal direction of free surface.dF dt F  t(VF   )0(6)158The pressure of arc Pa is considered to be a Gaussian distribution [21]: F 2  2 p22r 2  2 p in VOF model, the surface pressure can be expressed as a continuous body forces(CSF) [22] in the vicinity of the interface:r 2  2 pp aP maxexp()exp()(7)Farcpa1 2 F   ( 21)(8)where Pmax is maximum arc pressure, r is effective radius of the arc pressure, σp is arc pressure distribution coefficient, F is VOF gradient, and ρ1 and ρ2 are the density of first and second phases respectively.

Surface tension Fst can be written as follows by CSF method:  FF st1 2   ( 2)1(9)where γ is surface tension coefficient, κ is curvature of free surface.Marangoni shear force acting on metal free surface is given by, M    TT n(10)where T is temperature, n is the vector tangential to the local free surface.The enthalpy- porosity model is employed to process solidification and melting of metals. The liquid fraction is used to implicit track the solid-liquid interface.

A momentum source Fms is added to the momentum equation for representing mushy zone,Fms2  (1 ) 3    ) (Amush(V V  p)(11)where Amush is mushy zone constant, β is the liquid volume fraction, ε is a small number (0.001) to prevent division by zero, Vp is the solid velocity due to the pulling of solidified material out of the domain.The effects of the plasma arc on the metal surface are represented by an empirical plasma drag forceformula [23] to account for the plasma drag effect on the droplet in this study.FdragC ds1 2 g2 w g2  D d ( 4)(12)where Cds is the drag coefficient for a sphere, ρg is the plasma gas density; wg velocity, and Dd is the droplet diameter which determined by wire feed rate vw, wire radius rw, and droplet transfer frequency fd [24],  w  dis the assumed plasma gasv w f2 r w1/3 )D d(6d(13)where ρw and ρd are the mass densities of the electrode and droplet respectively.The electromagnetic force Fmag in molten pool is assumed to be symmetrical Gaussian distribution.The calculations of current and electromagnetic force are given by Kumar [25].JzId 2  jexp(2 r d 2  j)(1z c)(14)JrId c 2[1 exp(2 r d 2  j)](15)159B F x I m  2 r[1 exp( J B z x r 2 r d 2  j 2  I d m 2 2   r 2 jz c)](1)exp((16)2 r d 2  j )[1 exp(2 r d 2  j)](1z x ) c r(17)F y J B  zy r 2 I d  m 2 2   r 2 jexp(2 r d 2  j )[1 exp(2 r d 2  j)](1z y ) c r(18)F zJ B  r2 I m 2 2   4 jr[1 exp(2 r d 2  j2 )] (1z c)(19)In the above equations, Jz and Jr are represents the current density in axial and radial direction, Bθ ismagnetic flux density which is induced by welding current.

I is welding current, d is the current density distribution factor, σ j is the effective radius of the arc, r is radius, c is the thickness of the work piece, and μm is permeability. Fx, Fy and Fz are components of electromagnetic force in x, y and z direction.The external longitudinal static magnetic field can lead to bimodal distribution of current density at molten pool surface [20]. In this paper, the excitation current is 1.1A, and the induction flux density Bz at metal pool surface is 0.014T (Measured by Gauss meter), comparing with the ref.[19], the changing of welding current density can be ignored. When the external longitudinal static magnetic field is applied, the electromagnetic force can be expressed as,' F x' F yF xF y((JJrrBzB z))x r y r(20)(21)' F zF z(22)c) Source item of energyThe arc heat Sarc and droplet heat Sdrop are the energy input of weld pool.

The Sarc can beexpressed as a ellipsoid heat source,S arcq arcx y z ( , , )  ) 3 3 (1 UI  d    c2e 3 x 2  q2e 3y 2  q2e 3 z 2 c2(23)qwhere U is welding voltage, I is welding current, c is ellipsoid radius of depth direction, η is efficiency of arc thermal, ηd is proportion of droplet energy, σq is horizontal axis length of the ellipsoid heat source.The droplet heat resource Sdrop is calculated by means of droplet temperature Td.

2. The coordinate system of simulation is locating at top surface of weld bead which is shown in Fig.1.Fig. 2. A schematic representation of GMAW CFD simulationThe domain IJKLMNOP is metal phase, and the other is air phase.

4. Fig.4 (a)~(g) and Fig.4 (h)~(n) are the simulated results under the 0A are shown in Fig. 4. Fig.4 (a)~(g) and Fig.4 (h)~(n) are the simulated results under the 0A respectively. and 1.1A excitation current respectively162163164165Fig.4. Metal transition andanalysis results Metal transition and metal pool dynamics analysis resultsFrom Fig. 4, it can be clearly see molten pool is mainly flowing around the z axis the rotational movement. well as smaller temperature gradient is obtained in the pool as a result of the rotational movement well as smaller temperature gradient is distribution. At 0.048s to 0.052s, droplet olten pool isotherms become asymmetric distribution.

At 0.048s to 0.052s, droplet olten pool isotherms become asymmetric distribution. At 0.048s to 0.052s, Simultaneously, the molten pool isotherms become asymmetric impingements acting on molten pool are After 0.052s, molten metal began to fall when the longitudinal magnetic field weld bead will higher than normal depositionclearly seen that while the longitudinal magnetic field uniform temperature distribution as around the z axis at z=-2mm section. More uniform temperature distribution asmagnetic field is applied, the flow ofting on molten pool are more pronounced when the longitudinal magnetic fieldmagnetic field is applied. at the rear of molten pool is higherolten metal began to fall, it can be seen that the molten metal at the rear of mis applied at 0.06s, which can predict that surface eposition.0.06s, which can predict that surface ripples on solidifiedFig.5. Tangential-velocityath2 at 0.06s. (a) Path 1; (b) Path 2.

Tangential-velocityath2 at 0.06s. (a) Path 1; (b) Path 2. velocity evolvement on Path1 and Path2 at 0.06s. (a) Path 1; (b) Path 2.Fig.5 represents the distributions ofrepresents the distributions of tangential-velocity on path 1 (z=-2mm, y=0, x(2mm, y=0, x(-10mm~10mm))and path 2 (z=-2mm, x=0, y(-4mm~4mm)) at 0.06s. tangential velocity while external longitudinal magnetic filed front and rear of the molten pool, the t (a)). The tangential-velocity on path 2 Fig.5 (b)).4mm~4mm)) at 0.06s. It can be seen that the moltenpool has a significantexternal longitudinal magnetic filed is applied. Because ofecause of the asymmetry betweenpool, the tangential-velocity is not equal on path 1 at front on path 2 is substantially equal between left and right offront and rear (shown in Fig.5 between left and right of the molten pool (shown in166It can also be predicted from Fig. 4 and Fig.

4 and Fig. 5 that, the tangential motion of molten metal will drivethe molten metal moving to the edge of the pool and reduce the height of single-bead as well as increase the width. Simultaneously, the smaller temperature gradient is an important factor for small residual stress and distortion in AAM parts [18].ExperimentsThe schematic representation of longitudinal magnetic field assisted AAM is shown in Fig.1, andthe schematic diagram of the experimental set-up are shown in Fig.6, the arrows represent the direction of the material flow or information flow.Fig.6. Schematic diagram of the experimental set-up.The experimental process parameters are shown in Table 4.

Schematic diagram of the experimental set-up.The experimental process parameters are shown in Table 4. The welding mode is DC pulse mode which the process parameters include peak voltage(Upeak), base current(Ibase), wire feed rate(Vwire), welding speed(Vweld), frequency(f) and peak pulse time(Tpulse).Table 4 Experimental process parametersUpeak(V) 30Ibase(A) Vwire(m/min) Vweld(mm/min)357.5530f(Hz) 230Tpluse (ms) 3.1Excitation current(A) 1.1According to ref.3, the ratio of the wire feed rate to the welding speed(RWFRTWS) is more than12.5 , the cross-section profile can be fitted with an arc curve. The RWFRTWS of current experiment is more than 14, so the optimal center distance of overlapping beads can be calculated by arc curve overlapping mode, the specific calculation methods and formulas can be found in ref.3.

Fig.7(a) is the cross section of single weld bead, the left bead is normal deposition bead and the right bead is magnetic field assisted deposition bead. Fig.7(b) and Fig.7(c) are the cross section of one layer and four beads overlapping of normal deposition and magnetic field assisted deposition respectively. The sectional dimensions of weld bead are shown in Table 5. It can be seen that the width of single-bead increased as well as the height of bead decreased significantly when longitudinal static magnetic field was applied. The experimental data are in accordance with the predicted results. The optimal overlapping distance can be calculated and obtained a better overlapping bead morphology [3], however, it is impossible to achieve the ideally at overlapped surface which has been evidenced by experimental study [28, 29].

Compared with the normal deposition overlapping beads, the magnetic field assisted deposition overlapping beads are more flat and uniform and beneficial to stacking deposition for next layer and improving the surface quality and dimensional accuracy of the AAM parts.167Table 5 The sectional dimensions of weld bead TableItemsNormal deposition Magnetic field assisted depositionSingle weld beadHeight(mm) 2.5Width(mm) 8210The width of four beads he width of four beads overlapping deposition(mm) overlapping 2227Fig.7. The cross section of single weld beadsection of single weld bead and overlapping beads.

The cross section of single weld beadsection of single weld bead and overlapping beads. (a) single weld bead deposition; (c) four beads overlapping of magnetic fieldsingle weld bead; (b) four beadsoverlapping of normal depositionmagnetic field assisted deposition.ConclusionsA new arc based additive manufacturing method assisted with external longitudinal static magnetic additive manufacturing method assisted with external longitudinal static magnetic additive manufacturing method assisted with external longitudinal static magnetic orresponding numerical simulation and experimental investigationfield is proposed, and the corresponding deposition and external longitudinal magnetic filed assisted deposition are performed deposition and external longitudinal magnetic filed assisted depositi results are gained:umerical simulation and experimental investigation for normalperformed.

The following main1) The external longitudinal static magnetic fieldhe external longitudinal static magnetic field can induce the tangential stirring force in molten pool and induce the tangential stirring force in molten pool and can reduce the temperature gradient indrive the molten metal moving to the edge of the pool, which can reduce the temperature drive the molten metal moving to the edge of the pool forming region.that the tangential stirring force can reduce the height of single-bead 2) The related experimental results accordance with the predicted results. as well as increase the width of single-bead.

as well as increase the width of single-bead. The experimental data are in accordance with the predicted results as well as increase the width of singleal results show that the tangential stirring force can reduce the height of single3) The external longitudinal magnetic fieldping are more flat and uniform surface quality and dimensional beneficial to stacking deposition for next layer and improving the surface quality and dimensionalmagnetic field assisted multi-beads overlapping are more flat andwhich is beneficial to stacking deposition accuracy of the AAM parts.References[1] Stenbacka N, Choquet I, Hurtig K. Review of arc efficiency values for gas tungsten arc welding[C]//IIW Stenbacka N, Choquet I, Hurtig K. Review of arc efficiency values for gas tungsten arc welding[C]//IIW Stenbacka N, Choquet I, Hurtig K. Review of arc efficiency values for gas tungsten arc welding[C]//IIW 20 April, 2012. 2012: 1-21.

Science and Technology of Welding and Joining, 2013, 18(1): 32-37.[25] Kumar A, DebRoy T. Calculation of three-dimensional electromagnetic force field during arc welding[J]. Journal of Applied Physics, 2003,94(2):1267.[26] Rao Z H, Zhou J, Liao S M, et al. Three-dimensional modeling of transport phenomena and their effect on the formation of ripples in gas metal arc welding[J]. Journal of Applied Physics, 2010,107(5):54905.[27] Hu J, Tsai H L. Heat and mass transfer in gas metal arc welding. Part I: The arc[J]. International Journal of Heat and Mass Transfer, 2007,50(5-6):833-846.[28] Suryakumar S, Karunakaran KP, Bernard A, Chandrasekhar U, Raghavender N, Sharma D. Weld bead modeling and process optimization in hybrid layered manufacturing. Comput-Aided Des, 2011;43:331–44.[29] Aiyiti W, Zhao W, Lu B, Tang Y. Investigation of the overlapping parameters of MPAW-based rapid prototyping. Rapid Prototyp J, 2006;12:165–72.170.

gh Thermal Conductivity Coatings via LENS for Thermal Management Applications Félix A. España, Vamsi Krishna Balla, Susmita Bose and Amit Bandyopadhyay W. M. Keck Biomedical Materials Research Laboratory School of Mechanical and Materials Engineering, Washington State University Abstract Surface modification has been used to improve wear resistance, corrosion resistance and thermal barrier properties of metals. However, no significant attempts have been made to improve thermal conductivity by surface modification. In this work, we have examined the feasibility of enhancing thermal conductivity (TC) of stainless steel by depositing brass using Laser Engineered Net Shaping (LENS). The coating increased the TC of the substrate by 65% at 100 C°. Significantly low thermal contact resistance was observed between the coating and the substrate due to minimal dilution and defect free sound interface.

Significantly low thermal contact resistance was observed between the coating and the substrate due to minimal dilution and defect free sound interface. Our results indicate that laser processing can be used on low coefficient of thermal expansion metal matrix composites to create feature based coatings to enhance their heat transfer capability.Introduction Coating is a unique way to tailor the surface properties of a component to suit a specific environment without sacrificing the bulk characteristics of the structure. In recent years, lasers have been extensively used as a surface modification tool for engineering properties of materials and structures [1]. Laser can be used to develop a coating or cladding of a material having a property grossly different from that of the substrate. These coatings are metallurgically bonded, with the substrate providing a sound and adherent interface.

These coatings are metallurgically bonded, with the substrate providing a sound and adherent interface. The surface structure can be modified in accordance with the service requirement by varying process variables such as laser power, scan speed, beam diameter, precursor composition and precursor feed rate [2, 3]. Till date lasers have been used to deposit alloys and ceramics to create structures with high hardness, excellent wear [4] and erosion resistance [5], and highly encouraging thermal barrier properties [6, 7]. However, there have been no attempts to improve the thermal conductivity (TC) of a material via surface modification, except one preliminary study by our group [8]. Surface modification of an alloy substrate with high thermal conductivity coating can not only improve thermal properties but also thermal performance by providing faster heat dissipating surface layer.

Moreover, it has been shown that the laser-based deposition techniques usually ensure good metallurgical bond with uniform intermixed region between the coating and the substrate material [9]. Such structures with enhanced heat dissipating top layer can find potential applications in heat exchanges of automotive and power plant industries. For example, copper or copper alloy coating on stainless steel would provide necessary corrosion resistance on one side and high heat dissipating layer on the other side leading to overall enhanced thermal performance. Moreover, this coating approach can also be implemented on low coefficient of thermal expansion (CTE) metal matrix composites (MMCs), used in electronic components, to improve their thermal performance by providing enhanced heat dispersing surface layer. In line with this, the TC of Mo–Cu alloys has been increased, from ∼190 to about 245 W/m K, by bonding these alloy with outer copper layers (Cu/Mo–Cu/Cu) [10].

In line with this, the TC of Mo–Cu alloys has been increased, from ∼190 to about 245 W/m K, by bonding these alloy with outer copper layers (Cu/Mo–Cu/Cu) [10]. However, these laminates usually suffer from high interface thermal contact resistance due to sharp interface. On the other hand, laser-based deposition techniques usually ensure good metallurgical bond with uniform intermixed region between the coating and the substrate materials [9, 11]. In addition, laser 103–105 K/s, surface modification is characterized by rapid solidification rates, up to intermetallic compounds between eliminating metallurgically incompatible metals/alloys [12].

In addition, laser 103–105 K/s, surface modification is characterized by rapid solidification rates, up to intermetallic compounds between eliminating metallurgically incompatible metals/alloys [12]. This feature is extremely useful whenformation of undesirable equilibriumdeveloping a copper-based coating on steel base substrate, which has a high affinity to form cracks due to low solid solubilities, as in the present work.In this study, high thermal conductivity brass (70Cu-30Zn) coating was prepared on AISI 410 stainless steel substrate to enhance its inherent TC, using laser engineered net shaping (LENS™). The simplicity of the coating process and flexibility in selecting process parameters with a high degree of deposition accuracy for desired physical and mechanical properties at appropriate locations of the part are some of the advantages of LENS™.

This feature enables us to fabricate feature based high TC coating for enhanced thermal performance.Experimental procedure and Materials Commercial stainless steel grade 410 flat sheet of 3 mm thickness was used as the substrate material. Brass alloy powder (Sigma-Aldrich) with particle size range between 50-150µm, based on sieve analysis, was used to create the coating. The chemical composition of this particular alloy powder is copper-70wt% and zinc-30wt%. LENS™-750 (Optomec Inc. Albuquerque, NM, USA) fitted with a 0.5 kW Nd-YAG laser system was utilized to satisfactory manufacture the high thermal conductivity brass coating disks with 12.5 mm diameter. A laser power level of 400 W, i.e. corresponding to a laser beam energy density of 226 W/mm2, was used to partially melt the alloy powder during the deposition process to create the brass coating on the substrate. In this work, scan speed of 15 mm/s and powder feed rate of 22 g/min were employed.

In this work, scan speed of 15 mm/s and powder feed rate of 22 g/min were employed. The parameters were selected based on previous optimization to create fully dense, good-quality coatings i.e. no voids. The coating thickness was controlled by the number of layers deposited on the substrate. However, post-deposition sample polishing was done to achieve the desired coating thicknesses. The coating sample was segmented and mounted for metallographic analysis. The top surface along with the cross-section of the sample was polished, cleaned and then etched with a solution of FeCl3 (g), HCl (ml), and distilled water (ml) in a ratio of 1:20:100 to reveal the coating’s microstructure features. Microstructural characterization of the coating sample was evaluated using a field emission scanning electron microscope (FEI, 200F). Vicker’s microhardness measurements (Leco, M-400G3) were performed on the laser processed coating sample using 50 g load for 15 s.

Vicker’s microhardness measurements (Leco, M-400G3) were performed on the laser processed coating sample using 50 g load for 15 s. An average value of 5 measurements was reported.To evaluate the thermal properties of the substrate and the brass coating, the sample was polished and cleaned with alcohol, to remove any dirt particles and grease form the surface. The polished sample was lightly coated with graphite to ensure proper testing and accurate results, but mainly to ensure that the emissivity or absorptivity was high so that we can increase the energy absorbed on the flashed side of the sample and also to increase the temperature signal on the back side of the sample. The thermal diffusivity and specific heat were measured using the LFA 447 Nanoflash® (NETZSCH). The LFA 447 Nanoflash® is based on the well known flash method (ASTM E 1461-07) to obtain the thermal properties. This method uses a high-intensity short duration light energy pulse.

The LFA 447 Nanoflash® is based on the well known flash method (ASTM E 1461-07) to obtain the thermal properties. This method uses a high-intensity short duration light energy pulse. The energy of the pulse is absorbed on the front surface of the sample and the resulting rear face temperature rise is measured using an infrared detector. By analysis of the resulting temperature versus time curve, the thermal diffusivity can be determined. When the thermal diffusivity of the sample is to be determined over a temperature range, the measurement must be repeated at each temperature of interest. Simple geometry, small specimen size requirements, fast testing, easiness of handling and high accuracy makes the flash method very versatile and advantageous. Another advantage of the flash method is the factthat specific heat values can be obtained directly from the testing procedure using differential scanning calorimetry and a well know material as the reference.

The coating was tested at 25 to 250 °C. Each measurement taken was an average over 10 individual data points assuring accurate results.Results and Discussion Coating Microstructure and Hardness. Laser power level of 400 W with 15 mm/s scan speed and a powder feed rate of 22 g/min completely melted the brass powder resulting in a fully dense coating on the AISI 410 stainless steel. As shown in Figure 1a, defects such as cracks, residual porosity and voids were also absent in the coating, which could have detrimental effect on thermal properties of the coating/composites.(a)(b)Figure 1: (a) Microstructure showing defect free and fully dense brass coating on stainless steel substrate, (b) High magnification SEM microstructure of coating-substrate interface and composition variation across the interface.During laser processing the material’s laser absorptivity plays a special role, because it controls the energy input from the laser to the sample.

The laser absorption coefficient of the material is directly proportional to its electrical resistance and inversely proportional to the wavelength of the laser [13]. The relatively high electrical conductivity of brass than AISI 410 steel decreases its laser absorption coefficient. Under identical processing conditions of laser power and scan speed, the total energy absorbed by the brass will thus be lower than the energy absorbed by the AISI 410 steel substrate, which can lead to incomplete melting at the interface of individual layers and consequently poor interlayer bonding. However, in the present coating, complete melting between individual layers and between the coating and the substrate indicates that the laser energy input used to fabricate the coating was sufficient to completely melt the brass leading to fully dense product.Another important factor that is of great importance is the interface region between the brass coating and the substrate.

The interface was observed to be crack free and contained an intermixed region with uniform distribution of brass and steel throughout the region. A magnification SEM microstructure of the interface is shown in Figure 1b. During LENS™processing, initially the substrate is partially melted and then the powders are injected into the molten pool, this creates an intermixed/diffuse interface region rather than a sharp interface during plasma spraying or roll bonding to create such composites. The absence of sharp interface between the coating the substrate would reduce the thermal contact resistance at the interface and improves heat transfer capability between the two materials. The thermal contact resistance of the laser processed composite was measured between 25 and 250°C using the two layer multi- material mode available in LFA 447 Nanoflash®. It was observed that the temperature has no effect on the thermal contact resistance over the studied temperature range.

It was observed that the temperature has no effect on the thermal contact resistance over the studied temperature range. An average value of 1.982 × 10-11 m2K/W was recorded for the 0.45 mm brass coating on AISI 410 stainless steel. The extremely low contact resistance values were presumably arises due to the absence of sharp interface between the coating and the substrate. In addition, the concentration profile across the coating-substrate interface, shown in Figure 1b, illustrates that the laser deposition did not result in any sharp interface or any intermetallic formation within the coating sample. The EDS line scan also shows how the Cu-Zn infiltrates the AISI 410 stainless steel substrate more than the stainless steel on the Cu.

The EDS line scan also shows how the Cu-Zn infiltrates the AISI 410 stainless steel substrate more than the stainless steel on the Cu. This can be attributed to the melting of the substrate as we introduce the brass powder.Average hardness of the laser processed brass coating was 92 ± 6 HV and that of steel substrate was 360 ± 5 HV, while the average hardness of the intermixed region was around 146 HV. In actuality the hardness of the interface gradually increased from 100 HV to around 195 HV as the substrate was approached. Throughout the coating cross section no abrupt increase in the hardness was observed. The fact that we melt and then rapidly quench (fast cooling rate) does not allow enough time for the brass and stainless steel to form any intermetallics or cracks at the interface region. The same explanation holds for the low contact resistance measured from the experimental procedure for the high TC brass coating materials.

The same explanation holds for the low contact resistance measured from the experimental procedure for the high TC brass coating materials. There was a noticeable increase of hardness value of the substrate near the intermixed region. This was due to the melting and rapid solidification of the steel resulting in martensite in this steel. Laser surface melting has been shown to increase the hardness of AISI 410 stainless steel due to the formation of martensite during cooling [14]. The high hardness of substrate near the interface could also be due to solid solution strengthening of Cu in steel.Thermal Conductivity. The thermal conductivity, k, of the coating, and substrate was calculated , where α is the thermal diffusivity (m2/sec), cp is the using the following formula: specific heat (J/(kg⋅K), and ρ is the bulk density (kg/m3) of the material. The bulk density of the samples was calculated from sample geometry and mass.

The bulk density of the samples was calculated from sample geometry and mass. The thermal diffusivity and specific heat of the samples were experimentally determined. Figure 2 shows that the thermal conductivity increases with increasing temperature for the brass coated AISI 410 steel sample. However, the TC of AISI 410 substrate sample initially increased up to 100°C and remained unchanged up to 250°C. The TC values varied from 33 to 40 W/mK for the 0.45 mm thick brass coating and 20 to 23 W/mK for the SS410 substrate over the temperature range of 25 to 250°C. The TC of the substrate increased between 60 and 75% due to the deposition of the brass coating and test temperature. As experimentally determined specific heat and thermal diffusivity increased with increasing temperature for brass-steel composites the thermal conductivity of these samples also increased with temperature. The increase in the TC of the composites with the deposition of the brass coating is understandable.

The increase in the TC of the composites with the deposition of the brass coating is understandable. The ratio of coating thickness over the substrate thickness was 0.4. This high ratio means high volume of higher TC brass in the!"##=pckcomposites, enabling faster heat removal from the steel substrate and faster heat dissipation from the top surface to the environment.0.45mm thick coating experimentalAISI 410 substrate0.45mm thick coating theoretical1924293439050100150200250300Temperature (C)Thermal Conductivity (W/mK)Figure 2: Thermal conductivity of laser deposited high thermal conductivity brass coating both the theoretical and experimental values, and AISI 410 substrate.The main objective of the high TC brass coating on AISI 410 substrate is to enhance the heat dissipation in the direction perpendicular to the AISI 410 substrate surface.

The lower bound TC of multilayer materials (kML) can be calculated using a series heat-transfer model assuming slabgeometry and negligible interfacial resistance [15]:122121kVkVkkkML!+!=, whereVvolume fraction, k is the thermal conductivity, and the 1 and 2 subscripts refer to the individual material layers. As shown in Figure 2, the calculated multilayer TC values for the brass coating varied from 25 to 28 W/mK over the temperature range of 25 to 250 °C. The experimentally determined average TC of AISI 410 with brass coatings is higher than the theoretically estimated values, which supports the absence/extremely low interfacial thermal contact resistance in these laser processed composites. In multilayer materials, if the interfaces play a role in reducing heat conduction, then the experimentally determined TC values of AISI 410 steel with brass coating will be lower than those estimated using series heat-transfer model.

However, experimentally determined average TC of AISI 410 with the brass coatings is 32 to 42% higher than the theoretically estimated values. This supports the absence of interfacial thermal contact resistance in the present composite structures due to high degree of bonding, diffuse interface region and lack of additional porosity at the substrate–coating interface. The observed high TC could also be due to the supersaturated solid solution of Cu with steel near the substrate–coating interface,isthewhich would have eliminated any interfacial thermal resistance between brass and steel. It has been shown that the formation of supersaturated solid solutions during LENS™ processing is very common due to associated rapid cooling rates [16]. The results in this investigation are consistent with recent work on plasma-sprayed multilayer coatings [17], where the largest decrease in the TC of the coatings has been due to splat interfaces within each layer.

More efficient coatings can be fabricated via LENS™ by providing fins projecting out form the coating surface. Being a CAD based process LENS™ enables to create feature based high thermal conductivity coatings for enhanced thermal performance in various applications.Conclusions Laser surface modification has been successfully applied to show that by depositing high thermal conductivity material the thermal performance of substrate can be improved. The thermal conductivity and performance of AISI 410 stainless steel has been improved up to a range of 60 to 75%, due to defect free diffuse interface and dense coating structure.

The thermal conductivity and performance of AISI 410 stainless steel has been improved up to a range of 60 to 75%, due to defect free diffuse interface and dense coating structure. The LENS™ in combination with the coating approach can be employ to deposit other high thermal conductivity materials on different substrate material eliminating the interfacial problems associated with the sharp interface present in conventional processes.AcknowledgmentsAuthors acknowledge the financial support from National Science Foundation under theGrant number CMMI 0728348.References1. A. Agarwal and N. B. Dahotre, Int. J Refra. Met. & Hard Matls.17 [4] (1999) p. 283. 2. J Mazumder, “Laser assisted surface coatings”;, in Metallurgical and Ceramic Protective Coatings. Edited by K. H. Stern, Chapman & Hall, London, 1996, pp 74-111. 3. J. Feng, M. G. S. Ferreira and R. Vilar, Surf. and Coat. Technol. 88 [1-3] (1996) p. 212. 4.

GH VISCOSITY JETTING OF CONDUCTIVE AND DIELECTRIC PASTES FOR PRINTED ELECTRONICSJ. Ledesma-Fernandez, C. Tuck, R. HagueAdditive Manufacturing and 3D Printing Group, University of Nottingham, NG7 2RD, UK REVIEWEDAbstractInk-jet printing of multiple materials in 3 dimensions is a promising alternative to traditional patterning methods due to its flexibility, scalability and accuracy. However, the printability of the inks is strongly restricted by material properties such as surface tension and viscosity. Dispensing high viscosity fluids on a drop-on-demand approach is a potential solution that can facilitate the incorporation of new materials to the jetting catalogue. Consequently, in this study 2 micro-dispensing valves are used in combination with a mechanical stage to deposit conductive and dielectric pastes with viscosities of 15.3 ± 0.2 and 0.638 ± 0.005 Pa·s (at 25°C and 10 s-1 shear rate) respectively.

Crucial printing parameters such as pressure, temperature, pulse shape and drop spacing are studied in order to optimise the process. Additionally, post- printing characteristics such as contact angle of different materials and cured layer profiles are also measured and taken into account during the designing of the 3D patterns to minimise the negative effects of the thickness miss-match of different materials. Finally, the manufacturing capability of the set-up is demonstrated by the fabrication of a functional device using a combination of “pick-and-place” components and high viscosity jetting.IntroductionAdditive Manufacturing offers numerous advantages over traditional fabrication such as customizable parts, increased design freedom or complex parts at no extra cost. These are only a few of the characteristics that have contributed to the great expansion of Additive Manufacturing during the last decade.

These are only a few of the characteristics that have contributed to the great expansion of Additive Manufacturing during the last decade. However, most systems are only capable of fabricating objects with a single material at a time, which limits the functionality of the final part. Therefore, to fully take advantage of the capabilities of Additive Manufacturing the incorporation of multiple materials in a single process is required. This approach has already been explored for some additive technologies such as multi-nozzle FDM [1], inkjet-infiltrated porous parts built by Laser Sintering [2] and stereolithography performed with exchangeable resin vats [3].

However, although these approaches are a step forward in their respective fields, there are situations where other Additive Manufacturing techniques may be required.An alternative technology capable of multi-material fabrication is inkjet printing, which fundamental principle of depositing small drops of fluids in precise locations on demand is ideal for the creation of complex patterns with local variations of material concentration. This capability of creating “digital materials” by combining a finite number of primary inks was first commercialised by Objet Geometries Ltd. in early 2000 [4]. However, the capabilities of inkjet are not limited to the deposition of polymers and many other materials have been jetted in the last decades. Particularly, in the field of conductive materials inkjet printing has also been used to create functional transistors [5], strain sensors on fabric [6] or low-cost Radio-frequency40Identification (RFID) tags [7] among many other examples.

This has led to the rise of printed electronics, which is a potential solution for low-cost and customized products.However, inkjet printing imposes great restrictions to the properties of the inks that need to be finely formulated to avoid nozzle clogging, present long stability and behave according to predictive models during the printing process. One consequence of these boundaries is that the viscosity of the inks needs to remain typically under 25 mPa s, which limits the molecular weight of the used polymers and the solid content of particulate inks. This not only narrows considerably the material catalogue but also forces the addition of rheological modifiers such as solvents and surfactants [8] that alter the properties of the deposited tracks.

This not only narrows considerably the material catalogue but also forces the addition of rheological modifiers such as solvents and surfactants [8] that alter the properties of the deposited tracks. This is especially relevant in the case of conductive materials, where the organic elements added to improve the processability of the ink need to be removed during a post-processing step to allow the percolation and/or sintering of the conductive particles dispersed on the ink.A high viscosity non-contact approach would allow the printing of materials with higher solid content, which can translate into reduced volumetric shrinkage, increased mechanical properties and improved conductivity.

This has been explored with different approaches such as aerosol jet [9], where the material is atomized in a pressurized chamber previous deposition; acoustic inkjet, where an ultrasonic perturbation is focused on the fluid free surface to induce the creation of droplets without the need of a nozzle [10]; and ultrasonic droplet generation, where the nozzle plate is specifically designed to focus the acoustic waves generated by a piezoelectric actuator and produce the ejection of droplets [11]. Additionally, jetting valves that combine pneumatic pressure and mechanical actuation have also been used to create 2-dimensional patterns in a continuous inkjet approach [12]. In this work a similar jetting valve is used to deposit two fluids of high viscosity in a drop-on-demand approach to create 2D and 3D geometries.

In this work a similar jetting valve is used to deposit two fluids of high viscosity in a drop-on-demand approach to create 2D and 3D geometries. The different stages of drop formation, deposition and post-treatment of 2 materials are studied and their application for multi-material jetting is demonstrated by the creation of a proof-of-concept sample including embedded conductive tracks and an LED.Materials and MethodsHigh viscosity printing systemThe set-up consists of a computer controlled 3-dimensional stage with an overall accuracy of 5μm that controls the ejection and position of 2 PICO xMOD micro-dispensing valves (Nordson EFD). The system is equipped with a digital camera to monitor the printing stage and select the printing origin. A UV-lamp (FireFly 25x10, Phoseon Technology) was installed to enable photo-curing of materials during the process.

A UV-lamp (FireFly 25x10, Phoseon Technology) was installed to enable photo-curing of materials during the process. The software is based on a combination of C# and XML and is in charge of printing parameters such as printing speed, drop spacing and curing speed. This code also communicates with the digital controller (Nordson EFD) that provides the printing signal to the valves.The working principle of the jetting valves is based on the combination of pneumatic and mechanical actuation. Precisely, the fluid is pressurized (up to 0.6 MPa) and injected on the jetting chamber along a fluid path that can be heated up to 100° C. There, a piezoelectrically driven piston ended with a ceramic sealing ball is oscillated according to the printing signal (up to 500 Hz), opening and closing the nozzle as a consequence.

There, a piezoelectrically driven piston ended with a ceramic sealing ball is oscillated according to the printing signal (up to 500 Hz), opening and closing the nozzle as a consequence. These driving pulses are mainly41described by its pulse time (𝑡𝑝) and cycle time (𝑡𝑐), which are both under user control through manipulation of the digital controller, and the opening and closing speed of the valve, which is not under user control but can be selected from a limited amount of pre-sets. This mechanism allows the ejection of materials in the range of 50 – 200.000 mPa s [13]. A schematic illustration of this system is shown in Figure 1 together with a description of the structure of a train of printing pulses.Figure 1: (Left) Schematic illustration of the PICO valve with a description of its most relevant parts. Modified from valve user manual [13].

Modified from valve user manual [13]. (Right) Typical shape of the driving pulse showing the effect of the pulse, close and cycling times on the state of the valve [12].High viscosity printing systemThe viscosity of the materials was measured with a Kinexus Pro rheometer (Malvern Instruments Ltd.) with a cone-plate configuration. A shear rate of 1 – 50 s-1 was used for the carbon paste while a range of shear rates of 1-100 s-1 was applied to the photo-reactive resin. The range of temperatures was 25 – 40 °C, which is within reach of the micro-dispensing valve.The surface tension of the fluids was characterised with a Drop Shape Analyser (DSA100, Krüss) and a G18 stainless steel blunt needle, using the Pendant Drop Method and the Laplace-Young model provided by the built-in software (Drop Shape Analysis, DSA4).

Ten droplets of each fluid were measured at 21.5 ± 0.1 °C to obtain an average value ± standard error of the surface tension.The density of the fluid was calculated by depositing known volumes of material with a syringe of 0.1 ml precision and measuring the variation in weight using an analytical laboratory balance (SI-234 Denver Instrument) at an ambient temperature of 19.8 ± 0.1°C. A linear regression was later performed to obtain the density value ± standard error.Jettability and deposition studyThe drop footprints resulting from different parameter sets were studied with an optical microscope (Eclipse LV100 ND, Nikon) to study the jetting behaviour of high viscosity fluids.

Later, the jetting process related to selected parameter configurations were recorded with a high speed camera FASTCAM APX RS (Photron Inc.) illuminated by a collimated light source (OSL2 Fiber Illuminator, Thorlabs).42Post-process investigation and performance testUV-curable materials were post-processed with a UV-lamp of 2 W/cm2 peak irradiance and 380 – 420 nm output (FireFly 25x10, Phoseon Technology) at 4.0 mm ± 0.1 mm of the printing platform. Materials requiring evaporation of solvent between layers where left to dry under controlled circumstances.The study of the profile of multi-layered samples was performed with a surface profilometer (Surftest SV-3100, Mitutoyo), taking linear profiles at different points in the sample. Moreover, the profile of UV-curable multi-layered structures was also observed with an optical microscope (Eclipse LV100 ND, Nikon) and a sample holder that allowed the observation of the edge of the printed samples.

Their height was measured with an electronic outside micrometer (Schut Geometrical Technology) of 0.001 mm of precision 3 times to obtain the average value ± standard error.To calculate the resistivity of conductive materials several tracks were printed with different number of layers and left to dry at room temperature and pressure for 24 hours. Their resistance was measured 6 times using a digital multimeter (TENMA), accepting the average ± standard error as the resistance of that particular track. Then, their cross section was measured with a contact profilometer as previously explained.

Then, their cross section was measured with a contact profilometer as previously explained. Finally, the resistivity value was obtained from a linear fit of these results, taking into account the definition of resistance [14]:𝑅 = (𝑙 ∙ 𝜌) ∙1𝐴Where R is the resistance of the track, l its length, ρ is the resistivity of the material and Athe cross-section of the sample.Multi-material interaction and proofs-of-conceptThe contact angle of a material on different substrates was measured with a Drop Shape Analyser (DSA100, Krüss) equipped with a G18 stainless steel blunt needle, using the Sessile Drop mode and the fitting method of the “tangent”. Each measurement was taken after a minute of the drop ejection, which is a realistic time scale of the printing process.

Each measurement was taken after a minute of the drop ejection, which is a realistic time scale of the printing process. The substrates consisted on an untreated glass slide, a clean piece of Polycarbonate film (Makrofol, Bayer) and a glass slide coated with 40 µm of one of the printing materials using a K Hand Coater (RK PrintCoat Instruments Ltd.) and applying the corresponding post-treatment of each material.MaterialsThe conductive material used for the set of experiments presented in this work is a commercial carbon paste (Electric Paint, Bare Conductive). A transparent photo-reactive resin (Clear, Formlabs) originally intended for stereolithography was selected as a structural material. The embedded LED used in the embedded electronics proof-of-concept was a green PICOLED (P12 Series, ROHM) of 1.0 x 0.6 x 0.2 mm and was place manually with a pair of tweezers between printing cycles.

The embedded LED used in the embedded electronics proof-of-concept was a green PICOLED (P12 Series, ROHM) of 1.0 x 0.6 x 0.2 mm and was place manually with a pair of tweezers between printing cycles. Its position was checked with the mounted camera of the printing stage and it was powered using 3 standard AA batteries.43(1)Results and discussionMaterial characterizationThe viscosity of the carbon paste was studied at different shear rates and temperatures, as can be seen in Figure 2. An increase on the shear rate is observed to cause a noticeable decrease of the viscosity, which is known as “shear thinning” and is common in polymeric composites with solid filling [15]. The viscosity is also reduced when the material is heated, however this needs to be taken carefully since temperatures greater than 35°C have been observed to cause quick evaporation of the solvent on the nozzle, leading to blockages. Therefore this material was printed at 25°C.

Therefore this material was printed at 25°C. Similarly, the viscosity of the photo-reactive resin was studied (Figure 3) and the same trend of reduction in viscosity with increased temperature was observed. Nevertheless, the rheological behaviour of this material at increased shear rates is different than the conductive material, since the viscosity is maintained constant after a small initial reduction at shear rates below 3s-1. This behaviour is typical from Newtonian fluids but a measurement of the viscosity at a broader shear rate range is required in order to confirm this observation.3515.0) s a P (y t i s o c s i V3025201510) s a P (y t i s o c s i V14.013.012.0511.0010.0110 Shear Rate (s-1)100253035Temperature (⁰C)Figure 2: Viscosity of a carbon paste measured with a rheometer in a cone-plane configuration. (Left) Constant temperature (25°C) and shear rates between 1 and 50 s-1.

(Left) Constant temperature (25°C) and shear rates between 1 and 50 s-1. (Right) Constant shear rate (10 s-1) but different temperatures in the range of 25 to 40°C.0.700.7) s a P (y t i s o c s i V0.680.660.640.62) s a P (y t i s o c s i V0.60.50.40.30.20.10.600.0110 Shear Rate (s-1)1002530 Temperature (°C)35Figure 3: Viscosity of a photoactive resin measured with a rheometer in a cone-plane configuration. (Left) Constant temperature (25°C) and shear rates between 1 and 100 s-1. (Right) Constant shear rate but different temperatures in the range of 25 to 40°C.444040Additionally, the surface tension and density of these materials were characterised (Table 1).

(Right) Constant shear rate but different temperatures in the range of 25 to 40°C.444040Additionally, the surface tension and density of these materials were characterised (Table 1). These materials show viscosities above the threshold of 25 mPa s typically accepted for piezo based inkjet techniques and could not be jetted in a commercial inkjet printer.Table 1: Summary of relevant material properties of the fluids used in this workViscosity @ 10 s-1,25°C (Pa s) Surface Tension (mN/m)Carbon Paste 15.3 ± 0.257.8 ± 1.8Photo-reactive Resin 0.638 ± 0.00534.9 ± 0.2Density (g/ml)1.254 ± 0.0161.038 ± 0.012Jettability and Deposition studyIn this work, the distance between the nozzle and the substrate at any given time is referred as gap size. Figure 4 shows droplets of carbon paste jetted at different heights from the substrate.

Figure 4 shows droplets of carbon paste jetted at different heights from the substrate. It can be observed that gap sizes larger than 1 mm result in the splash of the droplet on impact, which negatively affects the printing resolution. Therefore, the gap size was kept at 0.8 mm for the rest of this work. Likewise, the behaviour of photo-reactive resin were studied at different gap size but no clear difference in the drop size or shape was observed in the rage of 0.5 to 3 mm, therefore the 0.8 mm gap was also used for this material.Figure 4: Individual drops of carbon paste impacting on a polycarbonate substrate at different distances from the nozzle imaged by reflexion microscospy.In a similar way than the gap size, different pressures and valve closing times were tested by printing arrays of 6 x 6 unconnected drops and studying their size and shape with a reflection microscope (Figure 5).

From these sets of images it can be concluded that the size and shape of the drops is not visibly affected by the pressure in the range of 0.1 to 0.5 MPa. A further investigation conducted with high speed videos of the printing process showed that the ejection only happens at the end of the pulse time, i.e. during the closing time of the valve. This means that the applied pressure is not enough to push the material through the nozzle even when it is open and its main role is to fill the nozzle chamber to prevent material stagnation, as it was previously observed by Yang et al. 2013 during their study of the ejection of fluids of different viscosities with a valve similar to the one used in this work. Alternatively, Figure 5 also shows that the closing time is clearly the driving factor in this viscosity range, which is also in agreement with the explanation of Yang et al. 2013.

Alternatively, Figure 5 also shows that the closing time is clearly the driving factor in this viscosity range, which is also in agreement with the explanation of Yang et al. 2013. It is observed that for the carbon paste the profile with a closing time of 400 µs does not provide enough kinetic energy to produce a clean ejection; instead, as it was observed on the high speed videos, the tail of the drop breaks after the landing of the drop, collapsing over the substrate in an uncontrolled way. On the contrary, the45profile with 120 µs produces bigger drops with good repeatability but still shows signs of splashing due to its excessive closing speed. The profile with a closing time of 200 µs provides smaller drops and good repeatability across different pressures for both materials, which will lead to a higher printing resolution.

The profile with a closing time of 200 µs provides smaller drops and good repeatability across different pressures for both materials, which will lead to a higher printing resolution. Conversely, less influence of the closing time is observed for the UV-curable resin, where circular drops are obtained for the profiles with closing times of 200 µs and 120 µs. However, this good printing performance is abruptly interrupted when the closing time goes over 200 µs, which was tested several times without success. Nevertheless, this was only attempted in the range of pressures tested in this experiment and it is not ruled out that a higher pressure can force the ejection with these slow closing profiles but an investigation of the printability at higher pressures was not in the scope of this work.

Consequently, the profile with 200 µs and the pressure of 0.1 MPa is selected for both materials with associated drop diameters of 591 ± 2 μm and 391 ± 2 μm for the photo-reactive resin and the carbon paste respectively.Figure 5: Optical study of drop arrays of (left) a carbon filled polymer and (right) a photo-reactive resin with the pressure of the valve changing for each column and a different valve closing time in each row.A general overview of the jetting and deposition stages was recorded at 18000 fps for each material with the printing parameters that have been selected along this report. In these videos is observed that the material travels from the nozzle to the substrate in one frame for the carbon paste (Figure 6) and two for the UV-curable resin (Figure 7), which can be used to provide a minimum boundary of the average speed of the drops of 14 m/s and 7 m/s respectively.

This estimation can be greatly improved analysing videos at a higher frame rate, which will be performed in future work. From Figure 7 it can also be seen how the greater speed of the carbon paste helps detach the jet tail that stays connected to the nozzle during almost 3 ms in the case of the photo-active resin. This need to be taken into account when the printing speed is selected, because if the head moves when the tail is not separated from the nozzle, non-circular droplets will be obtained due to the uncontrolled collapse of that body of fluid. A good example of this can be found in Figure 5, where the slowest closing speed of the valve created this effect on the carbon paste regardless the pressure. Furthermore, Figure 7 shows that even though the presence46of the tail may cause problems with the UV-curable resin, this material recombines faster and more homogeneously than the carbon paste, which can be appreciated at the final frames of each material.

Specifically, the carbon paste changes the shape of the material deposited but it is unable to form something close to a spherical cap under 3 ms, while the photo-resin would be almost done in less than 2 ms if the jetting tail were not present. This may also explain why the surface of the UV-curable material is always smoother than the carbon.Figure 6: Printing events for the carbon paste with a gap size of 0.8 mm. Videos recorded at 18.000 fpsand 4.06 µs of shutter speed.Figure 7: Printing events for the photo-reactive resin with a gap size of 0.8 mm. Videos recorded at 18.000 fps and 4.06 µs of shutter speed.Once the ejection and deposition of individual drops was studied, several square patterns with different values of DPI (drops per inch) were produced in order to find the drop spacing for each material that yields features with connected drops, no internal holes and edges as sharp as possible.

In the case of the carbon paste (Figure 8), unconnected drops are observed at small DPI and more defined edges are obtained as the drop space is reduced. However, instead of focusing47only on these 2-dimensional deformations it is important that we focus on the 3D aspect of these samples, understanding the DPI not only as a measurement of droplet spacing but also as magnitude directly related with the deposited volume on a given area. This way, it is possible to observe that the thickness and curvature of the surface of the squares are notably incremented at higher DPIs, which in the pictures can be perceived by the variation in the light reflection. This may seem a positive effect keeping in mind that the focus of this work is creating 3D structures, but these kind of curved surfaces will affect the wetting of the following layer, which in turn affect the next one and so forth, leading ultimately to domed geometries.

Accordingly, the DPI for the carbon paste was chosen as 100 for the rest of the experiments.The case of the photo-resistive resin (Figure 9) seems different because bulging of the squares surface is not apparent but a strong dependence between the DPI and the surface roughness is observed instead. It is important to remember that each material has been post- treated to induce solidification using a different approach. Particularly, the carbon paste solvent evaporation was accelerated by heating the printing platform at 35°C, while the resin was cured using a UV-lamp placed at 4mm and moving over the pattern one time at a curing speed of 150 mm/s. It is believed that precisely the curing step, even applied during a time as short as the one used in this experiment, may have caused shrinkage on the sample, causing the material on some edges to retract (as can be seen in Figure 9) and the middle to the sample to shrink in a repeatable pattern.

Nevertheless, this effect can be reduced controlling the amount of fluid deposited per unit of area, as can be observed for the sample with a corresponding DPI of 60, which does not have sharp edges but does not display a strong texture on its surface either. Therefore, from this moment onward the DPI for the clear resin will be limited to 65.Figure 8: Printed arrays of carbon paste droplets with different DPI. An increase in the height of the sample can be appreciated due to the change in the reflection of the surface of the high DPI squares.48Figure 9: Printed arrays of UV-curable droplets with different DPI. The size of the scale bar is 500 µm.

The size of the scale bar is 500 µm. An increase in the texture of the surface can be appreciated at DPIs greater than 70 but sharp edges are not formed at DPI lower than 60.Post-process investigation and multi-layer testsIn order to study the layer thickness of the carbon paste and the influence of the drying time in the printing process, 8 tracks of an increasing number of layers from 1 to 8 were printed leaving 0 s, 30 s and 300 s between layers. These drying times, however, refer to the time programmed to stop the printing process completely on the modified software and would have to be added to the time that the machine uses to clean and reposition the head between layers in order to obtain an absolute measurement of the time between the deposition of the last drop on one layer and the first drop of the next one.

The results are shown in Figure 10, where it can be observed that the optimum drying time that minimises the spreading of the paste and leads to the most significant growth in the z-direction is between 31 and 300 s, since both the 0 s and 30 s samples display bulged tracks as a clear effect of excessive fluid per layer. These conclusions however cannot be generalised since the drying time is strongly related with the area exposed to air, which in turn would depend on the particular pattern. Nonetheless, this test provides a good estimation for processing purposes and will be repeated in the future for different post-treatment methods to compare with the non-accelerated evaporation here shown.0 s30 s300 sFigure 10: Carbon multi-layered samples printed with a different waiting time between layers to study the effect of the evaporation time on the geometry of the tracks. The first tracks on the left have only one layer but an additional one is added in each track up to 8 layers.

The first tracks on the left have only one layer but an additional one is added in each track up to 8 layers. Later, these samples were also used to calculate the resistivity of the material, which explains the marks that the multimeter probes left on some of the samples.49After 24h of drying at room temperature, the sample printed with a programmed drying time of 300s was analysed with a surface profilometer (Surftest SV-3100, Mitutoyo) to obtain the values for layer thickness shown in Figure 11. A value of 53.6 ± 0.6 µm for the thickness of each layer was obtained applying a linear fit to the height of the tracks and it will be used as the increment in the Z-axis per printed layer from now on. Additionally, comparing the measured resistance of each track with the average cross-sectional area provided by the contact profilometer, the resistivity of the carbon paste was calculated.

Additionally, comparing the measured resistance of each track with the average cross-sectional area provided by the contact profilometer, the resistivity of the carbon paste was calculated. Figure 11 also shows the mentioned plot of the resistance against the reciprocal of the area of each track, applying equation (1), from the slope of its linear fit a value of (1.09 ± 0.03) ∙ 10−2 Ω 𝑚 was calculated for the resistivity of the carbon paste. This result is comparable to the reported volume resistivity for carbon black conductive filament for fuse deposition modelling (9 − 12 ∙ 10−2 Ω 𝑚 )[16]. However, the measured value is still six orders of magnitude bigger than the resistivity of bulk silver (1.587 ∙ 10−8 Ω 𝑚 at 20°C) [14].

However, the measured value is still six orders of magnitude bigger than the resistivity of bulk silver (1.587 ∙ 10−8 Ω 𝑚 at 20°C) [14]. Nonetheless, this is a limitation strictly based on the used material and not the set-up, which potentially can be used to deposit more conductive materials.)m μ (t h g i e H500400300200100y = 53.6x + 0.9 R² = 0.9993)Ω( e c n a t s i s e R1600 1400 1200 1000 800 600 400 200 0y = 69.8x + 48 R² = 0.99770024Number of layers680510 1/A (nm-2)15Figure 11: (Left) Linearity of the layer thickness on multi-layered carbon samples with 300 s of drying time per layer. (Right) Resistivity calculation of the carbon paste based on the comparison between their individual resistance and the cross-sectional information obtained from contact profilometry.The profile of multi-layered samples of photo-reactive resin was studied by printing samples of 30 layers with different curing speeds, i.e.

the speed of the UV-lamp over the printed pattern, to check the influence on the final geometry of this parameter. The profile of the structures was observed with a reflection microscope (Figure 12) and their height was measured with an electronic micrometer (Table 2). These results indicate that a curing speed of 150 mm/s at a distance of 4 mm from the sample is enough to cure the material and prevent it to flow before the deposition of the following layer, leading to analogous results than slower curing speeds. This speed could be potentially increased but it will not reduce significantly the length of the process since the longest stage is the printing, which is limited to a maximum of 500 Hz. Additionally, assuming linearity of the increase of the sample height with the number of layers it is possible to estimate the layer thickness of this material (Table 2) in order to use it in the multi-material samples printed in the last section of this work.

These multi-layered samples also revealed another interesting behavior of the UV-curable resin, since in the 3 cases there are some dark lines crossing the sample perpendicular to the substrate with a fairly constant frequency. A closer study revealed that the distance between these vertical patterns is 400 ± 20 µm. This value is in5020agreement with the drop spacing corresponding with a DPI of 65, which is 391 µm. Therefore, it can be assumed that this pattern is created due to the air bubbles produced and trapped due to the impact of the resin during the drop deposition stage. However, if we look back to Figure 9, this pattern is not appreciated in the single layer there studied, which leads us to believe that the roughness of the previous layer is in fact the driving factor of this effect.Figure 12: Cuboids of photoactive resin with 30 layers. Each structure was cured at a different speed, therefore limiting the radiation dose of the sample.

Each structure was cured at a different speed, therefore limiting the radiation dose of the sample. The used curing speeds were a) 150 mm/s, b) 30 mm/s and c) 6 mm/s. The profile was investigated via optical imaging.Table 2: Height and estimated layer thickness corresponding to structures of photo-resin with 30 layers UV- cured at different speeds.Curing Speed (± 1 mm/s) 150 30 6Height (± 0.007 mm) 1.415 1.403 1.398Layer thickness (± 0.2 µm) 47.2 46.8 46.6Multi-material tests and proof-of-conceptThe previous tests were performed for each material individually but in a real multi- material printing situation this would not describe the process accurately. Therefore, tests involving various materials need to be carried out to assess their combined behaviour. In the present work the focus is placed on the different wettability of each material on different substrates, which can lead to variations to the layer thickness than what was previously observed.

In the present work the focus is placed on the different wettability of each material on different substrates, which can lead to variations to the layer thickness than what was previously observed. Consequently, the contact angle of droplets deposited on different substrates was tested and presented in Figure 13. Here it can be observed that the UV-curable resin wets more than the carbon paste in all situations, explaining its bigger footprints. Moreover, there is a clear51difference between the case of carbon on a carbon substrate and the UV-curable material deposited on itself, which clarifies why the structures printed with carbon presented sharper edges than the ones created with the resin. However, the conclusions extracted from this study can only describe trends and should not be extrapolated to the printing situation straight away.

However, the conclusions extracted from this study can only describe trends and should not be extrapolated to the printing situation straight away. This is mainly due to the difference on the roughness of hand coated and printed substrates but also due to the great difference between the volume of the droplets used in this study (in the range of microliters) and the ones involved in printing (nanoliters).Figure 13: Spreading behaviour of the 3 materials used in this work in different combinations. (Left) Contact angle of deposited drops using a Drop Analysing System. (Right) Top image of samples printed with the micro- dispensing valves used in this work. The substrates were also printed and post-treated accordingly.Following this reasoning, another matrix of material combinations was created using the printing set-up to assess the variation of the drop diameter instead of its contact angle, as can also be found in Figure 13.

This experiment confirmed that the carbon paste wets all the surfaces less than the photo-reactive resin, leading to smaller drops. The sample of carbon paste on a carbon substrate displayed individual drops on the edge of the sample but none in its middle, which seems to be an indication that the substrate material may not have been totally dry on the centre and facilitate the spreading of the drops. Nevertheless, the drops that were left are smaller than the ones printed directly on the polycarbonate film, as it was anticipated by the contact angle in these materials (Figure 13). Additionally, a contradiction of the previous observations regarding the spreading of the carbon paste on a photo-resin substrate was found, since the drops of carbon are the smallest on this substrate.

This may be related, as previously discussed, with the difference in roughness of the substrate and the substantial variation in droplet volume.52On the other hand, the behaviour of the UV-curable resin matches with the wettability previously observed, where the drops are bigger on the polycarbonate film and slightly smaller on the carbon substrate. In the case of the photoactive resin on a substrate of the same material there are no droplets to observe, which is a clear indication of the coalescence of the drops and the good wetting already observed previously. It is expected that a similar test with a smaller DPI will lead to a situation where drops could be observed before merging.

It is expected that a similar test with a smaller DPI will lead to a situation where drops could be observed before merging. This situation does not hinder the printing of multi-layered samples of photo-reactive resin, as it was previously shown in Figure 12, since the outer edge of the pattern is preserved long enough to pin it in place during the curing step.Once both the conductive and structural materials were characterised and tested together without relevant interaction between them, a proof-of-concept sample was created to show the capabilities of the high viscosity jetting approach. To do it, the structure was divided in different “structural blocks” containing a different number of layers of each material in order to compensate their different thickness, as is described in Table 3. However, since there is an LED embedded on the sample, the thickness of the component need to be compensated as well.

However, since there is an LED embedded on the sample, the thickness of the component need to be compensated as well. This resulted into the functional proof-of-concept sample shown in Figure 14, where the LED is emitting a clearly visible green light while is powered by 4.5 V.Table 3: Description of the main “structural blocks” of a proof-of-concept sample for embedded electronics. The carbon paste acts in this case as a functional material due to its conductivity and the UV-curable resin as a structural component. An LED is manually introduced after the carbon tracks are printed.PatternMaterialNumber of layersFunctional (carbon paste)5Structural (UV-curable resin)10,7,4,5Figure 14: Embedded LED and conductive tracks proof-of-concept. a) First structural block of structural material. b) Second structural block of structural material. c) First 2 structural blocks, conductive block and LED. d) Totally embedded LED sample.

a) First structural block of structural material. b) Second structural block of structural material. c) First 2 structural blocks, conductive block and LED. d) Totally embedded LED sample. e) Two probes with a voltage of 4.5 V are used to power the LED, which emits a green light.53ConclusionsThis investigation studied the suitability of a high viscosity jetting valve based on the combination of pneumatic and mechanic forces as an Additive Manufacturing technique for multi-material 3-dimensional parts. Two fluids of different rheological behaviour were jetted and deposited in 2D and 3D geometries, identifying their difference in layer thickness as the main challenge for multi-material samples. This issue was minimised by the printing of “structural blocks” of equivalent height, using the values obtained during the characterization of multi- layered tracks.

This issue was minimised by the printing of “structural blocks” of equivalent height, using the values obtained during the characterization of multi- layered tracks. Additionally, the transparency and electronic conductivity of the materials investigated were applied to the creation of a proof-of-concept sample with embedded conductive tracks and an LED. Further investigation of the influence of the material properties on the final geometry and performance of the parts is required in order to achieve the full potential of this promising approach.AcknowledgmentThis research is supported by the EPSRC Centre for Innovative Manufacturing in Additive Manufacturing. Additionally, the authors are grateful to Ehab Saleh, Mark East, Mark Hardy and Joseph White for their contributions in different aspects of this research.References[1] D. Espalin, J. A. Ramirez, F. Medina, and R. Wicker, “Multi-material, multi-technologyFDM: exploring build process variations,” Rapid Prototyp.

GH VISCOSITY JETTING SYSTEM FOR 3D REACTIVE INKJET PRINTINGH. Yang, Y. He, C. Tuck, R. Wildman, I. Ashcroft, P. Dickens, R. HagueAdditive Manufacturing and 3D printing group, University of Nottingham, NG7 2RD, UKAbstractHigh viscosity ink is a potential solution for the improvement of current 3D ink jetting technology. In this study, experiments are carried out to study a series of inks with differing viscosities jetted with PicoDotTM jet valves under different process parameters of temperature and pressure. Results show that a wide range of ink materials from low viscosity (water like materials) to very high viscosity (thixotropic materials) can be jetted with the piezoelectric actuated jet valves without the need of solvents and surfactants. The jetting volume can be controlled under certain conditions between 2nl and 15nl.

The jetting volume can be controlled under certain conditions between 2nl and 15nl. The jetting performances for low, medium and high viscosity inks have been recorded by high-speed video photography.IntroductionIn the last two decades ink-jet printing technology has seen enormous development as a precision micro dispensing tool [1-3]. Traditionally inkjet printing has been mostly applied in the fabrication of 2D patterns or structures due to the requirement of low viscosity of inks. An amount of solvent is needed to dilute the printed materials to facilitate printing while the solvent also needs to evaporate quickly after printing. If not carefully controlled, the solvent can induce problems such as ink clotting or air pollution, which usually demands additional complex and corrective cleaning processes. To circumvent such a problem, in some applications, instead of using a solvent, low viscosity molecules are chosen as the ink for inkjet printing directly.

To circumvent such a problem, in some applications, instead of using a solvent, low viscosity molecules are chosen as the ink for inkjet printing directly. After printing, the droplet will quickly solidify through reaction under the initiation of a catalyst, heat or high energy radiation such as ultra-violet (UV) light. Such printing is known as reactive inkjet printing.With no solvent or very few solvents used in the printing process, problems identified in the traditional inkjet printing can be effectively avoided. Because of 100% solidification, the reactive inkjet printing is also promising for 3D printing. Using UV curable polymers, 3D inkjet printers have been successfully developed and commercialized such as those sold by Objet [4].

Using UV curable polymers, 3D inkjet printers have been successfully developed and commercialized such as those sold by Objet [4]. The majority of UV curable polymers used for 3D printing are low molecular weight acrylates and the mechanical properties are therefore limited in their industrial applications.To prompt new applications for Additive Manufacturing (AM) and 3D printing, the development of new reactive printing materials and new printing processes are significant issues for exploration in the development of reactive inkjet printing. Researchers have been working on different reactive monomers, oligomers or prepolymers, like UV curable epoxy [5], caprolactam [6, 7], two components polyurethane [8], anionic and cationic polyelectrolytes [9] and hydrogels [10]. Some of the traditional polymerization and processing technologies can be applied into RIJP with the reaction scale being reduced from the macro to the micro, such as crosslinking or using a reactive prepolymer.

An ink jettingsystem that can deal with wide range of ink viscosity one key to the success of such 3D printing.Viscosities of inks for traditional inkjet system is normally in the range of about 1- 40cP. Rosen et al.[11] and Kiyama et al [12] have developed ultrasonic droplet generation systems which are able to print fluids up to 3000 cP of viscosity. However, those jetting systems haven’t been used in printing-based additive manufacturing process due to the complexity and inconsistency of high viscosity jetting. This research will use a piezoelectric actuated jetting system, PicoDotTM jet valves[13], to jet materials of a wide range of viscosity. This jetting system combines mechanical and pneumatic mechanisms so that a higher force can be applied to the jetting materials. With the piezoelectric actuator, a high jetting frequency can be achieved for a fast manufacturing speed.

With the piezoelectric actuator, a high jetting frequency can be achieved for a fast manufacturing speed. In this study, the jetting capability of this system is evaluated and the possibility of using this system in 3D reactive inkjet printing is investigated.Materials and MethodsInkjet printing systemEFD PicoDotTM jet valves (Nordson EFD, UK) are used to jet inks under different temperatures and air pressures. Figure 1 shows the schematic drawing of the inside view of the jet valve and the assembled jetting system. Fluid fills the flow path and the jetting reservoir with a constant pressure from a syringe connected with compressed air through regulator, with the air pressure adjustable to up to 80psi. Driven by two piezoelectric actuators, a rod with a sealing ball made of wear-resistant ceramic at the lower end is lifted or lowered. The nozzle is closed by the sealing ball. When the sealing ball is lifted, fluid can flow through the valve.

The nozzle is closed by the sealing ball. When the sealing ball is lifted, fluid can flow through the valve. The extremely fast piezoelectric actuator helps to achieve a jetting frequency of up to 500 Hz. Depending on the fluid being using, the piezoelectric actuator can produce ink droplets of different size, with the smallest deposits being 2nl.A controller is connected with the jetting valve to control the jetting time and jetting temperature. The high resolution PicoDot valve controller (0.01 ms) makes it possible to adjust the deposit size precisely. The temperature of the fluid path is also controlled via the PicoDot valve controller. The valves feature a heated fluid path with a maximum temperature of +100°C. The controller can also be connected to a computer so that the jetting process can be automatically controlled by the computer.

The controller can also be connected to a computer so that the jetting process can be automatically controlled by the computer. A 3D stage used to enable 3D printing.(a) (b)Figure 1 Schematic of the PicoDotTM jetting valve (a) and the 3D jetting system (b)Figure 2 shows the assembly and the jetting nozzle. The jetting valve is fixed on a computer-controlled XYZ stages with in-house made fixture. Two jetting valves can be fixed at the same time for multi-material jetting.Figure 2 3D jetting systemJet Value OperationFigure 3 shows the working principle of the PicoDot dispersion systems of Nordson EFD jetting system. Regular signal generated by an amplifier located in the PicoDot jet valve controller works on the piezoelectric actuators to generate mechanical force to open/close the valve. One droplet was jetted from the orifice for each valve opening-closing cycle.

One droplet was jetted from the orifice for each valve opening-closing cycle. The minimum jetting time (one open/close cycle) at maximum stroke is approximately 0.120 milliseconds for PicoDot LV jet valves and 0.250 milliseconds for PicoDot MV jet valves.Figure 3 Signals generated by the jet valve controllerFor different materials, the droplet volume (V) can be controlled by optimizing the jetting parameters. The droplet volume is decided by two factors: the mechanical force and the pneumatic force (Equation 1). The material jetted due to the mechanical force is mainly decided by the structure of the jet valve, and the jetting rate (Qm) can be controlled by changing the close time (tclose). The jetting rate due to pneumatic force (Qp) can be calculated with Equation 2, which is decided by the power law constants n and K of a fluid, the pressure difference ΔP, and the nozzle configuration (diameter D and length L).

Overall the jetting volume with the pneumatic force is the jetting rate (Qp) times the pulse time (tp).QV mtclosetQ  ppQP  D  323     4 n 3n 1PD      LK 4    1 nMaterialsFive polymer materials with different viscosity were chosen to evaluate the jetting capability of the jetting valves. The rheology property of the materials was measured with a parallel plate rheometer (Kinexus, Malvern instruments, UK). Table 1 shows the rheological properties of the inks used.Table 1 Material ViscositiesJetting process studyThe PicoDotTM valves are available in several different configurations to handle fluids with the viscosity ranging from 50cP to 500,000 cP. Nozzles are available with 50 µm, 100 µm, 150µm, 200µm, 300µm and 400µm orifice diameters to accommodate a wide range of ink dot sizes.

Nozzles are available with 50 µm, 100 µm, 150µm, 200µm, 300µm and 400µm orifice diameters to accommodate a wide range of ink dot sizes. An LV type valve is suitable for the viscosity range from low viscosity up to 1000cP, and MV type valve for viscosity range from 1000cP up to 500,000cP, both with a 50 µm nozzle being used for the jetting evaluation. Several parameters need to be optimized to achieve suitable jetting quality and a suitable droplet volume, which includes the pulse time (valve opening time), jetting temperature and air pressure. For different materials jetted under different jetting parameters, the jetting volumes were measured by jetting 500 droplets and weighed with a 0.1mg analytical balance (Denver Instrument, USA).

For different materials jetted under different jetting parameters, the jetting volumes were measured by jetting 500 droplets and weighed with a 0.1mg analytical balance (Denver Instrument, USA). A high speed camera (FASTCAM APX RS, Photron, USA) was used to record the jetting process with a recording speed of 30,000 fps.(1)(2)ResultsMaterial CharacterizationFive ink materials with different viscosities were chosen for the jetting capability evaluation. For all the inks, the viscosity deceases with increasing temperature. Figure 4 shows the viscosity change of ink 3 under different temperatures, which is typical for all other materials. Below 40°C, the viscosity decreases quickly with the increase of temperature, and the rate of decrease slows when the temperature is above 40°C. Ink 4 and 5 are non-Newtonian fluids which have a shear thinning property. Figure 4 (b) shows the relationship of the viscosity and the shear rate.

Ink 4 and 5 are non-Newtonian fluids which have a shear thinning property. Figure 4 (b) shows the relationship of the viscosity and the shear rate. From this measurement, the power law constants n and K can be estimated. For ink 4, n is approximately 0.931 and for ink 5, n is only 0.041 and K is as high as 997. With increasing shear rate, ink 5 changes from solid to fluid, this makes it possible to be jetted through small orifices.700600s a P m, y t i s o c s i V5004003002001004432191209358010305070Temperature, °CFigure 4 (a) Viscosities under different temperature, ink 3. (b) Viscosity under different shear rate, ink 5, 25°CJetting parameters controlThe jetting parameters for the materials with different viscosities were investigated. Figure 5 shows the jetting parameters and jetting volumes for the low viscosity material (ink 1, 20cP at 25°C) and the high viscosity material (ink 5).

Figure 5 shows the jetting parameters and jetting volumes for the low viscosity material (ink 1, 20cP at 25°C) and the high viscosity material (ink 5). The LV type jetting valve was used for jetting ink 1 and MV jetting valve was used for ink 5. The results show that for low viscosity materials like ink 1and ink 2, the jetting volume increases with the increase of both the pulse time and the air pressure. This is because for low viscosity liquid, the pneumatic force is the determinant of the jetting volume. For high viscosity materials, the main function of the pneumatic force is to push the liquid to fill into the flow channel, and the droplet size is mainly decided by the mechanical force applied to the ink.

When the pulse time is less than a critical value, the jetting volume is mainly determined by the amount of the fluid filling the valve chamber, and the longer the pulse time, the more fluid is filled in the chamber; when the pulse time reaches the critical value, the filling process is saturated so the jetting volume will not change when the pulse time is further increasing.(a) (b)Figure 5 (a) Jetting parameters of Low viscosity material (Ink 1, 25°C, LV jetting valve). (b) Jetting performance of high viscosity material (Ink 5, 25°C, MV jetting valve)For a medium viscosity material like ink 3, it is possible to use both the LV type and the MV type jet valves. By controlling the jet valve temperature, the viscosity of ink 3 changes from 650cP (20°C) to 58cP (60°C). Table 2 shows the results of jetting ink 3 with LV jet valve under different temperatures. Higher temperatures reduce the fluid viscosity, thus the minimum pulse time and the jetting volume are reduced.Table 2.

Higher temperatures reduce the fluid viscosity, thus the minimum pulse time and the jetting volume are reduced.Table 2. Jetting parameters with LV jet valve (Ink3, 60Psi)Table 3. Jetting Parameters with MV jet valve (Ink 3, 25° (a), 20Psi (b))(a ) (b)Table 3 shows the results of using MV to jet ink 3. The results show that the jetting volumes didn’t change much for a certain range of pulse time (0.2 to 0.8ms), air pressure (20 to 60psi) and temperature (25 to 45°C). Pulse time as small as 0.2ms was used. Ink 4 with a medial viscosity of 4500cP also has similar jetting performance with the MV jet valve. This is because that the MV jet valve can generate much higher mechanical force than the LV jet valve, and the mechanical force is the main force for material jetting in this situation.Droplet formingThe process of material jetting, from fluid reaching the nozzle to droplet forming on a substrate, was recorded with a high speed camera.

The gap between the nozzle and the substrate is about 1.5mm. The jetting speed can be calculated with the displacement of the drop shown on the image divided by the time span. Figure 7(a) shows the jetting process of ink 2 using a LV jet valve. The low viscosity liquid (120cP, 25°C) was jetted with a speed of about 16m/s and a long tail is generated which can separated in less than 1ms. For the liquid with high viscosity like ink 4 (4500cP), the suitable jetting speed is about 6.8m/s and it takes up to a hundred milliseconds to get a round droplet on the substrate (Figure 7(b)). For the pseudoplastic material (ink 5) under a high jetting speed of about 20m/s, it can be jetted to form a droplet on the substrate with a tail connected to the nozzle. The droplet will keep the shape when jetted, until it is broken by the relative motion between the jetting valve and the substrate.

The droplet will keep the shape when jetted, until it is broken by the relative motion between the jetting valve and the substrate. Results show that even high viscosity pseudoplastic or thixotropic fluid can be jetted, using the MV jet valve, but the droplet may not be the ‘normal’ round shape.(a)(b)(c)Figure 7 Jetting process of (a) low viscosity, (b) medium and (c) high viscosity inksPatterningUse an XYZ stage, the jet valve location can be precisely controlled for planar feature printing, and the Z stage was controlled moving down vertically to match the printing speed for each deposited layer. Ideally (Figure 8a), the jet valve displacement speed should match the droplet forming speed, which is the drop size divided by the cycle time, to get a continuous pattern. Otherwise an unconnected pattern will be produced when the substrate motion is quicker than the droplet forming speed (Figure 8b).

Otherwise an unconnected pattern will be produced when the substrate motion is quicker than the droplet forming speed (Figure 8b). For the pseudoplastic fluid, a fiber rather than droplet will form, which needs extra attention on the jetting parameters to get an ideal jetting result (Figure 8c).Figure 8 (a) Pattering process, (b) droplets pattern and (c) jetted fiber of pseudoplastic fluidicConclusionReactive inkjet printing is a promising technology which may produce breakthrough for the functional materials used in 3D printing, although it is still in the preliminary stages. High viscosity jetting technique is one of the main issues to be resolved. This investigation evaluated the jetting capability of PicoDotTM jet valves for fluids of a wide range of viscosities. The parameters for jetting different viscosity fluids have been studied, which showed the high jetting flexibility of the jet valves.

The parameters for jetting different viscosity fluids have been studied, which showed the high jetting flexibility of the jet valves. Viscosities as low as 20cP and as high as 100,000cP (pseudoplastic) can all be precisely jetted. The droplets forming process has also been recorded with high speed camera, which provided useful information for further exploration of 3D patterning.AcknowledgementThis research is supported by the EPSRC Centre for Innovative Manufacturing inAdditive Manufacturing.References[1] M Singh, H M Haverinen, P Dhagat, et al., Inkjet Printing—Process and Its Applications, Adv. Mater. 2010, 22, 673–685.[2] C J Ferris, K J Gilmore, S Beirne, Bio-ink for on-demand printing of living cells, Biomater. Sci., 2013, 1, 224.[3] Takami Akagi, Tomoko Fujiwara, and Mitsuru Akashi, Rapid Fabrication of Polylactide Stereocomplex Using Layer-by-Layer Deposition by Inkjet Printing, Angew. Chem. Int. Ed.

Robinson3, K. Rupel4, R. Booth5, S. Kubiak61EOS GmbH, Krailling, Germany 2The Boeing Company, Everett, WA, USA 3formerly The Boeing Company, now 3DSIM, Park City, UT, USA 4The Boeing Company, St. Luis, MO, USA 5Advanced Laser Materials, Temple, TX, USA 6Stratasys Direct Manufacturing, Austin, TX, USAAbstractIn a time where rapid prototyping successively transforms to additive manufacturing (AM), nylon 11 and 12 and their composite powders, which have evolved to be the most commonly used materials in laser sintering (LS) due to their easy processability, cannot fulfil all challenging requirements of industrial applications any more. Especially in the aerospace industry, there is a high demand for stiff and lightweight parts for interiors, which currently are fabricated from glass fiber reinforced phenolic and epoxy resins by a lamination process.

Due to the strong diversity of the parts, this traditional manufacturing is quite labor- intensive and expensive, which makes it very attractive to manufacture these parts with additive manufacturing, especially laser sintering. Additional part design requirements, such as greater chemical and UV resistance, an elevated softening temperature, higher mechanical strength and better performance in flammability and heat release tests generate opportunities for the use of high performance AM polymers. Promising candidates that have the potential of satisfying these demands can be found among the different Polyaryletherketone thermoplastics. In this work we present the development of a carbon fiber filled PEKK composite material for laser sintering, optimized especially for the production of interiors, such as air ducts for cabin ventilation in aerospace application.

Based on process tests, powder characterization and test builds, the material and its manufacturing procedure were optimized towards isotropic properties and refreshability. Simulations of building cycles helped to understand the extent of powder ageing, which is directly connected to the ability to recycle the material. Furthermore the laser sintering hardware of an EOSINT P800 and the building processes were adapted to ensure a stable building process and fulfill the requirements of parts on mechanical properties in x, y and z directions, dimensional stability and surface quality.IntroductionEven in recent years, nylon powders are still the most commonly used polymers in powder bed fusion processes [Woh2016, p.54].

Looking at the laser sintering powder manufacturers, about 90% of the powders available in their 2015 portfolio are based on nylon 11 and 12 [Schm2015], and probably more than 90% of laser sintered parts worldwide were built from nylon 12 powders and their compounds [Schm2015, Schm2015.1]. Nylon powders are easy to process, are well known and understood. Many studies have focused on the mechanical properties, refreshability and characterization of different nylon powders manufactured for laser sintering [Good2012, Mun2015, VdE2015, Bai2015 and Ver2016]. Now, as additive manufacturing techniques evolve from rapid prototyping to series production, the demand for a more diversified spectrum of materials increases. Asking users of additive manufacturing about their requests for new materials, 60% would like to see high temperature plastics and 52% of the respondents desired carbon fiber reinforced materials, a request especially arising from automotive and aerospace industries [SDM2013].

In 2016, AM equipment and material manufacturers and AM service providers served many different industries; a significant share of AM parts (16.6%) goes to the aerospace industry [Woh2016, p.19f.]. This industry is known to have very demanding requirements on material properties, especially in terms of continuous operating temperature, flammability, smoke development and toxicity as well as chemical resistance. In addition, another 808very important property of parts used in aerospace applications is their weight, as it has direct influence on the fuel consumption of an aircraft. So far, many parts are manufactured from glass and carbon fiber reinforced phenolic and epoxy resins by a lamination process. This fabrication method is very labor-intensive and therefore expensive. Replacing these parts by additively manufactured ones has many advantages, including great freedom of design.

This fabrication method is very labor-intensive and therefore expensive. Replacing these parts by additively manufactured ones has many advantages, including great freedom of design. Here, we show the approach and first data on the development of carbon-fiber reinforced PEKK polymer, a high-performance polymer optimized for aerospace applications.Materials and EquipmentCarbon-fiber reinforced PEKK was supplied by ALM Inc., TX. The PEKK resin contains 23% carbon fibers and is sold under the name “HT23”. All Parts were built on an EOSINT P800, which was slightly modified to process carbon-fiber reinforced PEKK in an optimal way. The modifications include an optimized process, heating parameters and recoating concept as well as the use of a different recoating blade.

The modifications include an optimized process, heating parameters and recoating concept as well as the use of a different recoating blade. All parts were built after thoroughly warming up the machine employing an automated heat-up phase (about 3 hours), the building procedure was followed by a regulated cool-down phase [Pfi2009] to ensure uniform cooldown of all parts and avoid warpage. Melt viscosity was determined according to ISO 1133-1:2011 and recorded as melt volume rate [cm³/10min]. The lower the MVR value, the higher the melt viscosity of the examined material. Mechanical data was achieved from standard ASTM D638 tensile bars which were built in 3 spatial directions, i.e. an x-tensile bar was built with its gauge length parallel to the recoating direction (x-direction, cf. figure 1).Figure 1: Position of an x-tensile bar on the EOSINT P800 building platform.Approach and ResultsAmong industry are Polyaryletherketone polymers.

figure 1).Figure 1: Position of an x-tensile bar on the EOSINT P800 building platform.Approach and ResultsAmong industry are Polyaryletherketone polymers. These semi-crystalline polymers have been shown to be inherently flame retardant, develop minimal toxic gases when burned and have excellent chemical resistance. In addition, the continuous operating temperature is stated in literature to lie above 250°C for most of the polymers [Becker/Braun, p.359ff; Fink p.209ff.]. PEKK was chosen as a base polymer, a polymer which is commercially and readily available. Milled carbon fibers act as reinforcing component.the most promising candidatesto fulfilthe requirements ofthe aerospaceIsotropy: The most common way of reinforcing materials for laser sintering with carbon fibers, glass beads or other additives is to simply blend the polymer powder with the reinforcing component [Wan2015, Wan2015.1 and Ath2010].

This method has, especially for the use of carbon fibers, one major drawback: The part properties of the laser sintered parts exhibit high anisotropy. Due to the movement of the recoater, fibers tend to align along the recoating direction. This leads to strongly improved mechanical part properties in direction of the recoater movement (x-direction), but a significant drop in all other lateral dimensions (y and especially z). One method 809to counteract this disadvantage is to encapsulate the fibers into the grains of the powder [Lyo2014]. This enables the homogeneous orientation of the fibers in all spatial directions.Figure 2: Micrograph of HT23 powder grains with encapsulated carbon fibers.Table 1 summarizes the relative mechanical part properties in y- and z-direction, compared to x-direction for laser sintered parts from three different LS materials.

OXFAB® ESD produced and employed by Oxford Performance Materials (OPM), a blend of PEKK with carbon fibers, shows a drop in mechanical properties of up to 37% [OPM2016] while the two powders HP11-30 (manufactured by ALM Inc.)[Boo2016] and HT23, both having the carbon fibers encapsulated inside the grain (see figure 2), show significantly improved isotropic mechanical properties, which is of great interest for part designers.UTS [rel] Young’s Modulus [rel]XYZOXFAB® ESD (OPM)1 10.63 0.70HP11-30 (ALM)1 1 1 1 0.82 0.79HT23 (ALM) (process not final) 1 1 1 1 0.85 0.90Table 1: UTS and Young's Modulus for all three spatial directions normalized to x-direction.Cross sections of laser sintered parts visualize the distribution of carbon fibers throughout the different spatial directions. Figure 3a) shows micrographs of three different cuts through a cuboid part, illustrating the homogeneous distribution of the fibers.

Figure 3a) shows micrographs of three different cuts through a cuboid part, illustrating the homogeneous distribution of the fibers. In contrary, the micrograph (cut along the xz-plane) of EOS HP3 dry- blended with 10% carbon fibers indicates that most fibers are oriented in-plane, while hardly any fibers are oriented perpendicular to the sintered layers (Figure 3b)810Figure 3: Micrographs of part cross sections. a) Cuts in three directions through a cuboid laser part made from HT23. b) Cross section of a laser sintered part made from a HP3/carbon fiber dry-blend. The recoating direction is indicated in red.Refreshability: In the laser sintering process, the powder bed is heated to a temperature close to the melting temperature in order to ensure a secure building process and avoid warpage.

After finishing a build, 80-90% of the powder remains unsintered but has been exposed to long heating times which cause physical and chemical changes, generally including an increase of the melt viscosity in the material, that limit it’s reusability [Pha2008; Dot2009; Ghi2013]. Therefore, the used powder is usually sieved and blended with a certain fraction of new powder. Especially with regard to cost and sustainability, reuse of the not sintered powder is strongly favored. The extent of the viscosity decrease is dependent on several material- and processing parameters, such as thermooxidative stabilizing of the polymer or the building temperature. While the former can be influenced by additives or the base polymer itself, the latter is defined or changed in the building process. In order to get a first idea about the refreshability of HT23, a simulation of nine building cycles with different refresh rates was performed.

In order to get a first idea about the refreshability of HT23, a simulation of nine building cycles with different refresh rates was performed. Small amounts of powder were aged in an oven at a temperature comparable to the usual process chamber temperature and an atmosphere similar to the one in the EOSINT P800. After every oven cycle the oven aged powder was blended with either 60% or 50% of new powder and aged again. A sample of every blend was taken in order to determine the melt volume rate. Both experiments (cf. Figure 4) show an initial drop of the MVR value when comparing new powder (oven cycle “0”) with the refreshed ones, indicating an increase of the melt viscosity. Interestingly, there is no further increase of melt viscosity after the first oven 811cycle; it seems to stay approximately constant during all cycles. Even though the values for oven cycles 1 to 9 vary, they’re still within the tolerance of the measurements (cf. error bars in fig.

Even though the values for oven cycles 1 to 9 vary, they’re still within the tolerance of the measurements (cf. error bars in fig. 4a and b).Figure 4: Simulation of 9 cycles of building and refreshing the powder. Cycle "0" indicates the melt volume rate (MVR) of new powder. After every cycle, the powder was blended with 60% (a) or 50% (b) new powder and aged again.Comparing different refresh rates, it is easily discernable that the initial increase of the melt viscosity is more pronounced when using 50% new powder (Figure 4b) compared to when blending with 60% new powder (Figure 4a). Nevertheless, after the first refreshing cycle, the viscosity stays pretty constant as well. A similar behavior is seen when comparing the evolution of the melt viscosity with the mechanical data. UTS and Young’s Modulus of tensile bars built in x-direction drop by approximately 1% when using refreshed powder (60% new powder) instead of new powder.

UTS and Young’s Modulus of tensile bars built in x-direction drop by approximately 1% when using refreshed powder (60% new powder) instead of new powder. The decrease in the other spatial directions is within the same range.Summary and ApplicationsIn this work, we show the approach for the development of a carbon fiber reinforced PEKK high performance polymer for laser sintering. The material is optimized for the demanding requirements of aerospace industry. By encapsulating the carbon fibers into the polymer grain, nearly isotropic part properties could be achieved in contrast to carbon fiber and polymer dry-blends which exhibit significant part property differences over the spatial directions. Cross-sections of laser sintered parts indicate the homogeneous and isotropic distribution of carbon fibers throughout the parts. Several building and refreshing cycles were simulated by aging the powder in an oven and refreshing it.

Several building and refreshing cycles were simulated by aging the powder in an oven and refreshing it. A sample was taken from every cycle before aging the powder again and the melt viscosity was determined. Two experiments with refreshing rates of 60% and 50% showed that, even if there is a slight increase in viscosity between new powder and refreshed powder, the viscosity is approximately constant from the first refreshing cycle. First laser sintering results confirm this behavior. Mechanical data from parts built with refreshed powder (60% new) are insignificantly lower than the ones from parts built with new powder. In addition to isotropic part properties and refreshability, the stiff and lightweight material – as known for PAEK polymers – shows a high continuous operating temperature, great chemical resistance and inherent flame retardancy.

Facilitated by meeting the FAR 25.853 (requirements on flammability, heat release …) HT23 is perfectly suitable for applications under rough environments with demanding requirements on health and safety. Possible applications include air ducts and exhausts as well as highly customized interior parts for aircrafts and other industries with similar standards.812References[Woh2016] [Schm2015] [Schm2015.1] M. Schmid, A. amado, K. Wegener: Polymer Powders for selective Laser Sintering (SLS),T. Wohler: Wohler’s Report 2016, Wohlers Associates, (2016). M. Schmid: Selektives Lasersintern (SLS) mit Kunststoffen, Hanser, München (2015).[Good2012][Mun2015][VdE2015][Bai2015][Ver2016][SDM2013][Pfi20009][Bec1994][Fin2008] [Wan2015][Wan2015.1][Ath2010][Lyo2014][OPM2016] [Boo2016][Pha2008][Dot2009]AIP Conf. Proc. 1664 (2015) 160009. R.D. Goodridge, C.J. Tuck, R.J.M.

To avoid the formation of defects, which affect mechanical properties, a huge amount of specific know-how is crucial. Especially for regulated industries, such as medical or aerospace, this is a limiting factor for the widespread usage. In this work, high-speed imaging in combination with a high-magnification optic is used to gain deeper insight into the property-determining mechanisms and boundary conditions during the process. Thereby, the intensity distribution from the melt pool radiation is measured and analyzed with an imaging script to determine width, length, and cooling rate with a resolution of 1.44 µm/pixel. The potential of this data for predicting resulting scan track properties is demonstrated. It can be shown that an automatic width measurement deviates from the manually measured value by only 1.2% and the length measurement by 1.4%.

It can be shown that an automatic width measurement deviates from the manually measured value by only 1.2% and the length measurement by 1.4%. It is also possible to detect anomalies in the process, such as balling effects.Keywords: Additive Manufacturing, PBF-LB/M, Process Monitoring, High Resolution, Balling EffectIntroductionIn the realm of advanced manufacturing technologies, powder bed fusion using a laser (PBF-LB) [1] has emerged as a game-changing process, revolutionizing the production of complex and customized metallic components. PBF-LB is an additive manufacturing technology that utilizes a high-powered laser beam to selectively fuse metal powder. The exposure is repeated layer by layer until a three-dimensional object is created. [2–4] Thereby, the process has a high potential for weight reduction and the production of customized workpieces with a batch size of one, making it particularly suitable for aerospace or medical applications [5, 6].

In view of small quantities, predictive and holistic quality assurance is required. However, the PBF-LB process is susceptible to various defects that can compromise the integrity and quality of the final product. To mitigate these challenges, the implementation of robust process monitoring systems has become imperative and more knowledge must be gained about the origin of such defects. This contribution demonstrates an approach to gain deeper knowledge about the defect emergence. Thereby, the used system is based on a high-speed camera with high magnification where the gained data is evaluated automatically with a Python script. Overall, the method should initially help to detect geometrical or cooling anomalies of the melt within individual scan tracks.1116During the PBF-LB/M process, specific irregularities of the melt track as balling or even specific defects within the bulk material, like lack of fusion (LOF), gas pores or thermal stress- induced cracking can arise.

All listed phenomena and voids in the bulk material can lead to unstable processes or even structural weaknesses and reduced mechanical properties in the manufactured components [7–13]. Thus, it is needed to understand the emergence of these defects. In many studies, the occurrence of defects and their shape is investigated in dependence on the used process parameters like the laser power 𝑃𝐿, scan velocity 𝑣𝑆 or layer height 𝑑 [14–16]. Figure 1 provides an overview of the process parameters investigated by Gordon et al. [13]. Also shown are the volumetric data obtained by micro computed tomography (mCT), and the derived thresholds indicating the occurrence of gas pores or LOF defects. As in the other studies, Gorden et al.

As in the other studies, Gorden et al. found that LOF occurs when the set volumetric energy input is reduced, while gas pores occur when the applied volumetric energy is increased.Figure 1 –mCT data of PBF-LB/M manufactured specimen over the used process parameters – boundaries for the occurrence of defect types of LOF and gas pores [14].Although optimal process parameters are determined for material processing, implying a defect- free fabrication of parts, localized defects can still manifest. This can be attributed to various influencing factors that increase or decrease the energy input into the material volume. An example of reduced energy input and thus the occurrence of LOF is the interaction of the laser beam with process by-products, such as metal vapor or metal particles in solid or liquid form, which are stirred up during the process. [17, 18] Ladewig et al. [18] stated that even balling effects can occur depending on the laser-metal vapor interaction.

[17, 18] Ladewig et al. [18] stated that even balling effects can occur depending on the laser-metal vapor interaction. The balling behavior of 316L stainless steel processed on a self-developed SLM machine with a 200 W laser was investigated in dependence of serval factors by Liu et al. [19]. Thereby, among other things, they stated, that the behavior is dependent on two factors, the set laser power 𝑃𝐿 but also on the layer height 𝑑 respectively on the amount of material that must be melted. The influence of parameter variations on single tracks are shown in Figure 2a) and b). In Figure 2c) the consequences of balling at a local point within one layer are shown, where melt tracks do not overlap and thus gaps arise. These then can lead to defects like LOF. Based on the Figure 2, it can be considered that the melt pool geometry changes1117if the process conditions vary.

These then can lead to defects like LOF. Based on the Figure 2, it can be considered that the melt pool geometry changes1117if the process conditions vary. Since the effect of balling has a direct influence on the resulting material properties and the process stability, the occurrence of balling should be avoided.Figure 2 - Balling behavior of 316L stainless steel a) in dependence of set laser power 𝑷𝑳; b) layer height 𝒅 (both from [19]); c) Expression of balling within a layer [20].However, it is not possible to avoid the occurrence of balling totally. To minimize the production of parts with these defects, it makes sense to develop monitoring systems that can detect changes in the melt pool geometries during the process and take corrective action. As Grasso et al. [21] reviewed, there are many different approaches to monitor the PBF- LB/M process in situ, based on various methods. Thereby, five levels of process monitoring are introduced, as shown in Figure 3.

[21] reviewed, there are many different approaches to monitor the PBF- LB/M process in situ, based on various methods. Thereby, five levels of process monitoring are introduced, as shown in Figure 3. Level 0 offers lowest resolution of local phenomena, and level 4 is defines measurements, with the highest level of detail and thus the melt pool dynamics. In their work, they also stated that most of the publications about in situ measurement systems referred to systems in level 2 with over 34% followed by level 3 with 26% in the years from 2018 - 2020. The high percentage (51%) of publications looking at the process in Levels 2 and 3 shows that there is still a need to gain more knowledge there. Goossens et al. [22] have developed an approach for analyzing melt pool widths using data captured by a coaxially aligned high-speed camera.

Goossens et al. [22] have developed an approach for analyzing melt pool widths using data captured by a coaxially aligned high-speed camera. During the experiments, an EoSens® 3CL high-speed camera from Mikrotron GmbH was used to generate videos with a frame rate of 20,000 FPS, corresponding to a shutter time of 50 µs, and a spatial resolution of 11.8 µm x 11.8 µm per pixel. It has been demonstrated that the widths of melt pools, produced using varying combinations of process parameters (scan speed ranging from 400 mm/s to 1100 mm/s; laser power ranging from 100 W to 600 W), can be determined with an average error of -2.8% and a standard deviation of 18.5%. However, it has also been shown that melt pool widths can be underestimated by up to -54.05% or overestimated by up to 35.59%, depending on the parameter combination. The authors also acknowledge the coarse resolution as a potentialthethehighest.

The authors also acknowledge the coarse resolution as a potentialthethehighest. Level3Figure 3 - Different levels of measurements in the PBF- LB/M process according to Grasso et al. [21].1118source of error and refer to the ratio between the spatial resolution of a pixel and the expected melt pool width, which is 10.32%.In this contribution, a system is developed and applied which can be classified as level 3 due to its high temporal (40,000 frames per second acquisition rate) and spatial (1.44 µm x 1.44 µm per pixel) resolution to overcome high standard deviations. According to Grasso et al.

According to Grasso et al. [21] and recent studies, the spatial resolution, particularly when combined with the temporal resolution, corresponds to the highest resolution that has been used so far for investigating the melt pool dynamics or geometry in PBF-LB/M.MethodsEvaluation approachTo understand the underlying mechanisms in the defect formation, a method to investigate the melt pool morphology of individual scan vectors in the PBF-LB/M process was developed. For this purpose, a high-speed camera system (HSC system) is employed to record in situ, temporally, and spatially high-resolution videos during the exposures in PBF-LB/M. Subsequently, this data is analyzed regarding the width, length, and cooling time of the melt tracks with a self-developed Python script.

Subsequently, this data is analyzed regarding the width, length, and cooling time of the melt tracks with a self-developed Python script. The hardware and software used as well as the basics of the calculation are presented below.Test BenchThe experimental setup used to analyze the melt pool dynamics on a PBF-LB/M system with a high-speed camera (HSC) system is depicted in Figure 4. It illustrates all the key components along with their respective positions.Figure 4 - Schematic representation of the experimental setup for recording the melt pool dynamics using a high-speed camera. All investigations presented in this paper are carried out on an SLM 280 Hl 1.0 PBF-LB/M System manufactured and distributed by SLM Solutions AG. The system is equipped with two independent lasers which operate at a wavelength of 1,064 nm. Each laser is capable of emitting a maximum power of 700 W.

The system is equipped with two independent lasers which operate at a wavelength of 1,064 nm. Each laser is capable of emitting a maximum power of 700 W. The laser focal point diameter 𝑑𝐿 can be adjusted between 80 µm and 115 µm and is set to 90 µm for these experiments. Nitrogen is used as the process gas to evacuate the1119oxygen content during the processes as much as possible, but at least below 0.1 % to minimize the influence of oxidation. The gas flow is directed in negative x-direction (see Figure 4). [23]For the melt pool analysis, a Nova Fastcam S6 High-Speed Camera Ⓐ, manufactured by Photron GmbH, is utilized. It is mounted horizontally on a bracket (not depicted in Figure 4) that is equipped with a enables shifting 20.48 mm x 20.48 mm large image sensor, which resolves at 1,024 pixels x 1,024 pixels, resulting in a size of 20 µm x 20 µm of each. The monochrome image sensor offers a high light sensitivity of ISO 64,000.

The monochrome image sensor offers a high light sensitivity of ISO 64,000. The maximum frame rate of the camera at the highest resolution is 6,400 frames per second (FPS) but can be increased up to 800,000 FPS by reducing the utilized number of pixels. In the monochrome version of the camera, the intensity of light hitting the sensor is outputted in values ranging from 0 for a black pixel to 255 for a white pixel. If the detectable intensity falls below or exceeds these values, the respective pixel will output a value of 0 or 255. An overview of the resolution in relation to the maximum FPS and the resulting shutter speed is provided in Tablein a small rangein y-direction. The camera1. [24]Table 1 – Maximum usable resolution and shutter speed of the HSP system in dependence on the chosen FPSNear-IR SP785 Dichroic Shortpass Filter Ⓑ, provided by Midwest Optical Systems Inc., is utilized to reduce stray radiations during the analysis intensities.

The of melt pool in specified wavelength range which radiation is transmitted with over 90 % efficiency is between 425 nm and 770 nm. Less than 0.02 % of the intensity emitted by the SLM 280 HL laser is transmitted through this filter. [25]AnResolution 1,024 x 1,024 1,024 x 336 896 x 176 640 x 112Shutter speed in µs 154.2 48 23 11.1FPS in 1/s 6,400 20,000 40,000 80,000A 12X Distance Macro Zoom Lens, manufactured by Navitar Inc., is used as the objective Ⓔ. The combination of camera and objective enables magnification and thus high- resolution imaging of the field of view (FOV) with 𝑟𝐻𝑆𝐶 = 1.44 µm/pixel. The working distance 𝑎𝑊 of the entire system, which is the sum of the value 𝑎 and 𝑏 (see Figure 4), is adjustable from 73 mm to 85 mm. [26]Due to the small 𝑎𝑊 and thus height 𝑏, it is not possible to use the original recoater of the SLM 280 system.

[26]Due to the small 𝑎𝑊 and thus height 𝑏, it is not possible to use the original recoater of the SLM 280 system. For this purpose, a customized recoater Ⓕ with a lower height was developed to enable recoating within the process.Since the camera and the objective are aligned horizontally, a mirror Ⓖ is needed to redirect the radiation path. Therefore, a silver-coated quartz glass substrate with a surface flatness of λ/20 was used. The mirror reflects incident spectral radiation with wavelengths ranging from 450 nm to 10,000 nm with over 98% efficiency. [27]Since the lens is positioned inside the build chamber Ⓝ, the point where it enters the build chamber through the door Ⓓ of the SLM 280 HL had to be sealed. This is to ensure that the process conditions do not deviate from those of other processes.

This is to ensure that the process conditions do not deviate from those of other processes. For this purpose, a seal was designed and manufactured using the MEX process - shown in Ⓒ.Approach to evaluationThe evaluation of the melt pool widths 𝑏𝑆 and lengths 𝑙𝑆 is based on Planck's radiation law, which states that black bodies radiate with the power 𝑃𝐼 depending on their own temperature 𝑇. 0(𝜆, 𝑇), of a black body can be determined based on Equation 1 after The spectral radiance, 𝑀𝜆 wavelength representation. [28–30]1120𝟎(𝝀, 𝑻)𝒅𝑨𝒅𝝀 =𝑴𝝀𝟐𝝅𝒉𝒄𝟐 𝝀𝟓𝟏 𝒆𝒉𝒄/𝝀𝒌𝑻 − 𝟏𝒅𝑨𝒅𝝀 [𝑾 𝒎𝟑]𝝀 Wavelength 𝒉 𝒌 Boltzmann constant 𝒄 𝑻 Temperature 𝑨 0(𝜆, 𝑇) thereby represents the emitted radiant power of a defined area A of a black body within 𝑀𝜆 a wavelength range 𝑑𝜆. For a gray body or a selective radiator, which neither absorbs nor emits the 0(𝜆, 𝑇) must be corrected by multiplication with a factor, the total incident radiation, the 𝑀𝜆 emissivity 𝜀.

For a gray body or a selective radiator, which neither absorbs nor emits the 0(𝜆, 𝑇) must be corrected by multiplication with a factor, the total incident radiation, the 𝑀𝜆 emissivity 𝜀. The observed material, EOS 316L [31], does not have the properties of a black body, so the value of 𝜀 is needed to correlate the temperature with the intensities. Since the determination of 𝜀 is complicated caused by the temperature- and wavelength dependence, a direct correlation of the intensities detected by the camera and the real temperature is not carried out in this study. However, it is possible to distinguish between solid and molten material in the HSC videos and thus clearly identify the point at which the material property changes. When this point is detected, the intensity can be correlated to the melting or solidification temperature of the material. From there, the width and length of the melt pool can be determined.

When this point is detected, the intensity can be correlated to the melting or solidification temperature of the material. From there, the width and length of the melt pool can be determined. Since this method is based on the assumption, that the detected intensity is just caused by the radiation of the melt pool, it is important to evaluate and minimize the influence of other radiations or reflections. Thereby, the reflection of the PBF-LB/M laser must be considered. In Figure 5 the wavelength-dependent relative spectral response of the camera sensor and the transmission of the filter are plotted. To be able to compare these functions, both are plotted as the value effective intensity 𝐼𝐸, where this value describes howPlanck constant Speed of light Area𝐽𝑠 𝑚/𝑠 𝑚𝑚2𝑛𝑚 𝐽/𝐾 𝐾Figure 5 - Effective intensities the camera sensor, the SP785 filter and the whole HSC-system, taken from [24, 25].1121(1)much of the emitted intensity of a specimen can be detected.

For an evaluation of the total 𝐼𝐸 of the HSC system both functions 𝐼𝐸_𝑆𝑒𝑛𝑠𝑜𝑟 and 𝐼𝐸_𝑆𝑃785 are multiplied. The result is given as the red line, where it can be shown that the intensity of the PBF-LB/M laser can be detected by the cameras sensor with just 2.5𝑥10−4 %. Thus, the influence of the laser for the evaluation of the melt pool geometry can be considered as negligible. Similarly, the use of filters can prevent damage to the camera's sensor.Data analysis with Python scriptTo automate the analysis of the HSC data, a Python-based image analysis script was developed. This section describes the functionality of this tool. Overall, the script allows the determination of the following properties:1. Melt pool width 𝑏𝑆. 2. Melt pool length 𝑙𝑆. 3. Duration of liquidus phase Δ𝜏.Before presenting the evaluation methodology, an overview of the database used is given. Therefore, Figure 6a) shows an image from a video taken in this study that was evaluated with the script.

Therefore, Figure 6a) shows an image from a video taken in this study that was evaluated with the script. During the data acquisition for the analysis of the melt pool geometry, no external illumination is applied to the FOV in order not to influence the radiation effects of the melt pool. The scan vector in the x-direction, which indicates the laser exposure direction, is marked, as well as the coordinate system which is identical to that of the PBF-LB/M system, as shown in Figure 6a). Based on the coordinate system, the extent of 𝑏𝑆 is defined by 𝛥𝑦, and 𝑙𝑆 is defined by 𝛥𝑥. Therefore, it is necessary to consistently record the videos in the same orientation for such anFigure 6 - Images recorded during exposure of a single scan vector, a) without illumination as used for the experiments; b) with illumination to illustrate the melt path widths to be measured.evaluation.

Additionally, in Figure 6b) a frame is shown which was taken while the FOV was illuminated with an external light source. This allows the resulting melt pool width and length to be determined by manual measurements, which results then are used to calibrate the Python script.1122In the first step, the videos processed by the Python script are analyzed in terms of the occurring intensities 𝐼𝑥,𝑦. For this purpose, the intensity is analyzed for each pixel 𝑃𝑥,𝑦(𝜏) for each frame 1 ≤ 𝜏 ≤ 𝜏𝑚𝑎𝑥 and stored in a matrix. This matrix serves as the basis for evaluating the selected properties.The melt pool width 𝒃𝑺(𝒙) in µm is analyzed for each column of pixels 1 ≤ 𝑥 ≤ 𝑥𝑚𝑎𝑥. Therefore, the parameter 𝐼𝑚𝑒𝑙 is introduced. 𝐼𝑚𝑒𝑙 represents a material-specific intensity above which it is assumed that the material is in a molten state.

Therefore, the parameter 𝐼𝑚𝑒𝑙 is introduced. 𝐼𝑚𝑒𝑙 represents a material-specific intensity above which it is assumed that the material is in a molten state. For the determination of 𝐼𝑚𝑒𝑙 a single track is generated using the standard process parameters for processing EOS 316L on an SLM 280 HL machine while the HSC system is used to gain a dataset of the process. After exposure of the vector, an image of the generated welding track is taken while the FOV is illuminated with a light source, comparable with Figure 6b). Thereby, it must be ensured, that the melt track is clearly visible. Based on this image, 𝑛𝑚 measurements of the melt pool width are executed, and with these, the mean value 𝑏𝑟𝑒𝑎𝑙 and standard deviation 𝜎𝑟𝑒𝑎𝑙 are determined. Then, the obtained dataset of the HSC system is evaluated using the Python script, iteratively substituting a value for 𝐼𝑚𝑒𝑙. After each evaluation, the detected 𝑏𝑠 is compared with the 𝑏𝑟𝑒𝑎𝑙.

Then, the obtained dataset of the HSC system is evaluated using the Python script, iteratively substituting a value for 𝐼𝑚𝑒𝑙. After each evaluation, the detected 𝑏𝑠 is compared with the 𝑏𝑟𝑒𝑎𝑙. Once the deviation between 𝑏𝑠 and 𝑏𝑟𝑒𝑎𝑙 has reached the lowest possible value, 𝐼𝑚𝑒𝑙 is applied.To determine the 𝑏𝑠, the number of connected pixels ∆𝑦(𝑥, 𝜏) in the y-direction is analyzed in each column 𝑥 and for each frame 𝜏, that exceed the defined value of 𝐼𝑚𝑒𝑙. Then, the highest value of ∆𝑦𝑥𝑖, described as Δy𝑥𝑖,𝑚𝑎𝑥, which was analyzed in all analyzed frames at each x-position is used to determine the 𝑏𝑠 according to Formula (2).𝒃𝑺,𝒙𝒊(𝒙) = 𝚫𝐲𝒙𝒊,𝒎𝒂𝒙(𝒙, 𝝉) ∙ 𝒓𝑯𝑺𝑪Thereby, 𝑟𝐻𝑆𝐶 denotes the resolution of the camera system, with 𝑟𝐻𝑆𝐶 = 1.44 µm/pixel. An example for a determination of a melt pool width at one x-position is given in section ‘Results’.To determine the length of the melt pool 𝒍𝑺 in µm, the evolution of 𝐼𝑥,𝑦𝑆 (𝜏) along a line profile, in this case, the scan vector, is examined.

The essential condition for evaluating the melt pool length is similar to that of 𝑏𝑆, where the intensity must initially satisfy the condition 𝐼𝑥,𝑦𝑆 (𝜏) ≥ 𝐼𝑚𝑒𝑙. Subsequently, the number of frames, 𝛥𝜏𝑥𝑖,𝑦𝑆, is determined until the intensity value falls below an intensity where the material is in the solid phase again. The calculation of 𝑙𝑆 is then performed using Equation 3.𝒍𝑺,𝒚𝑺 =𝚫𝝉𝒚𝑺𝒗𝒔 𝑭𝑷𝑺Thereby, Δ𝜏𝑦𝑖 is multiplied by the scan velocity 𝑣𝑠 and divided by the applied frame rate 𝐹𝑃𝑆. As a result, it represents the calculated melt pool length of each pixel along the scan vector. To validate the results of the determination of the 𝑙𝑆 and thus correct choice of 𝐼𝑚𝑒𝑙 videos are taken and the datasets analyzed by the Python script. Furthermore, high-speed videos are recorded, where the same external light source as for the investigation of 𝑏𝑆 is used to illuminate to FOV. With these videos, the 𝑙𝑟𝑒𝑎𝑙 can be determined manually by observing the solidification points of the melt pool.

With these videos, the 𝑙𝑟𝑒𝑎𝑙 can be determined manually by observing the solidification points of the melt pool. Examples of this method are given in section ‘Results’.In addition to determining the melt pool lengths, 𝚫𝝉𝒙𝒊,𝒚𝒊 can also be used to assess the duration of the material being in a molten state at a defined position 𝑃𝑥,𝑦. Evaluating this time span, which is given in frames, can provide insights into local anomalies. Therefore, this value is1123(2)(3)determined for each pixel in the dataset as described before and plotted in graphs which indicate different Δ𝜏𝑥𝑖,𝑦𝑖 with a color scale.Measurement MethodsFor the investigations in this paper, four parameter combinations are utilized for the PBF-LB/M process, all of which are presented in Table 2. Parameter set 2, which is the standard parameter set for the manufacturing of 316L, is used to determine the value of 𝐼 where the material melts, as described before.

Parameter set 2, which is the standard parameter set for the manufacturing of 316L, is used to determine the value of 𝐼 where the material melts, as described before. Parameter sets 1 through 4 are then used to validate the results of the script-based melt pool width analysis. Thereby, the determined 𝐼 is used to evaluate the data with the Python script. The width of the scan tracks are also evaluated with microscopes for comparison.Table 2 - Used process parameters of the PBF- and HSC-systemParameter 𝑃𝐿 𝑣𝑆 ℎ 𝑑 𝑑𝐿 Material Process gas Oxygen level FPS Shutter speed Pixel used (x,y) ResolutionSet 1 170 WSet 2 200 WSet 3 260 WSet 4 320 Ws s e c o r P M B L800 mm/s 0.1 mm 0.03 mm 0.09 mm EOS 316L [31] Nitrogen ≤ 0.05 % 40,000 23 µs 896 x 176 1.44 µm x 1.44 µm per pixel/F B Pm e t s y SC S HParameter set 4 is applied in experiments where defects, specifically the balling effect, are provoked in single scan tracks.

The increased energy input, while keeping other conditions constant, forms the basis for these experiments.For all experiments carried out, the parameters to record the videos with the HSC system are the same. As described, a high magnification which results in a resolution of 1.44 µm x 1.44 µm per pixel is used while taking 40,000 frames per second. With this and the given scan velocity 𝑣𝑠 it can be determined that the movement of the laser is 20 µm per frame.1124ResultsCalibration of the width measurementTypical intensity profiles which arise if the intensity is determined along a line profile for consecutive frames from 𝑡1 = 0 µ𝑠 to 𝑡4 = 225 µ𝑠 are shown in Figure 7. Thereby, the profiles were obtained along the green line marked in the image, which is aligned in the y-direction with a constant value for 𝑥𝑖. It can be observed that the values of the intensity profiles decrease for consecutive frames.

It can be observed that the values of the intensity profiles decrease for consecutive frames. This effect can be attributed to the cooling of the melt pool at the given position with advancing time and distance between the evaluated line and the laser spot. The intensity threshold 𝐼𝑚𝑒𝑙, which must be calibrated, is represented by a red vertical line. If the measured 𝐼 exceeds this value, the material can be determined as in the liquid state, below as a solid. The value of 𝐼𝑚𝑒𝑙 = 149, whose calibration is described in the next paragraph, and can be used to determine the melt pool width 𝑏𝑆. Therefore, the furthest intersection points (∆𝑦𝑥𝑖,𝑚𝑎𝑥 = 𝑦2 − 𝑦1) between the intensity profiles and the threshold line are identified, as explained in section ‘Measurement Methods’. In this case, the 𝑏𝑆 = 114.98 µ𝑚 is determined by analyzing the intersections of the intensity profile in frame 𝑡1 = 0 µ𝑠.

In this case, the 𝑏𝑆 = 114.98 µ𝑚 is determined by analyzing the intersections of the intensity profile in frame 𝑡1 = 0 µ𝑠. Overall, it can be seen that an increase of the value of 𝐼𝑚𝑒𝑙 results in smaller determined melt pools, while a decrease of 𝐼𝑚𝑒𝑙 leads to an increase in the determined width. It is evident from the shape of the intensity profile that the intensity variations at the edges of the melt pool are approximately symmetric. For an accurate determination of 𝐼𝑚𝑒𝑙, it is not sufficient to examine only one 𝑥𝑖-position of the melt pool; instead, all determined widths for each x-position should be considered.Figure 7 - Intensities along a line profile directed in y-direction in different frames in the process.In Figure 8, all analyzed 𝑏𝑆(𝑥𝑖), which were determined using the calibrated value for 𝐼𝑚𝑒𝑙, are depicted as black dots. 𝐼𝑚𝑒𝑙 = 149 was determined as described in section ‘Measurement Methods’.

𝐼𝑚𝑒𝑙 = 149 was determined as described in section ‘Measurement Methods’. Therefore, the actual melt pool width 𝑏𝑟𝑒𝑎𝑙 and the standard deviation 𝜎𝑟𝑒𝑎𝑙, were determined to be 𝑏𝑟𝑒𝑎𝑙 = 119 µm ± 9 µm while using an image, taken with the camera while the1125FOV was illuminated, also according to section ‘Measurement Methods’. This image and the measurement lines are depicted in Figure 8. The mean value and standard deviation are represented by the red line and the red dashed lines, respectively. It can be observed that the widths determined by the script with 𝑏𝑆 = 117.12 µm ± 2.52 µm is within the range of the 𝑏𝑟𝑒𝑎𝑙 ± 𝜎𝑟𝑒𝑎𝑙. Overall, these results suggest that outliers in the melt pool widths are slightly underestimated to some extent by the Python script. Thereby, the lower 𝜎𝑆 can be attributed to the reduced error that occurs in this measurement.

Thereby, the lower 𝜎𝑆 can be attributed to the reduced error that occurs in this measurement. When manually measuring 𝑏𝑟𝑒𝑎𝑙, human errors such as subjective placement of the start and end points or slight shifts during positioning or variations in illumination must be considered. Also, the number of measurements carried out is much lower than in the Python-based analysis, so the influence of maxima and minima on the 𝑏𝑟𝑒𝑎𝑙 could be disproportionate. In this case, the deviation between 𝑏𝑠 and 𝑏𝑟𝑒𝑎𝑙 is quite low at 1.2%.

In this case, the deviation between 𝑏𝑠 and 𝑏𝑟𝑒𝑎𝑙 is quite low at 1.2%. Therefore, the calibration performed can be considered successful, as well as the approach to use the parameter 𝐼𝑚𝑒𝑙 to evaluate melt pool width.Figure 8 – Melt pool width determined by the Python script (𝟏 ≤ 𝒙𝒊 ≤ 𝟖𝟗𝟔), at the bottom a image where the melt pool width was measured manually (bottom).Validation of the width measurementsTo validate the functionality of the script-based evaluation of melt pool widths, new data sets are generated while exposure of single tracks using the four parameter sets presented in Table 2. The widths of the melt pools were initially determined using the Python script with the value for 𝐼 defined in the previous section. Subsequently, a manual analysis of the melt pool widths was conducted using a microscope for comparison. The results of these evaluations are shown in Figure 9 as 𝑏𝑆 and 𝑏𝑟𝑒𝑎𝑙, but also as frames of each evaluated data set.

The results of these evaluations are shown in Figure 9 as 𝑏𝑆 and 𝑏𝑟𝑒𝑎𝑙, but also as frames of each evaluated data set. The width corresponses as expected and in accordance with various studies, as described by Metelkova et al. [32] for example:1126The lowest laser power setting resulted in the smallest melt pool width, while the melt pool width increased with higher laser power. Furthermore, it is evident that the standard deviation increases as the selected 𝑃𝐿 deviates further from the standard parameter set for the processing of 316L due to the occurrence of variations in the melt track like balling. Even though 𝑏𝑟𝑒𝑎𝑙_𝑠𝑒𝑡2 = 116.99 ± 6.84 µm is slightly lower in this experiment compared to the result of the calibration measurement (𝑏𝑟𝑒𝑎𝑙_𝑐𝑎𝑙 = 119.12 ± 8.72 µm), the deviation between 𝑏𝑟𝑒𝑎𝑙 𝑠𝑒𝑡2 and 𝑏𝑆_𝑠𝑒𝑡2 = 115.73 ± 3.16 µm is only 1.09%, which is even lower than during calibration.

Similarly, it is observed that the discrepancies in varied parameter sets in these investigations are at most 1.72%, as seen in the width result of parameter set 4. Overall, the results of 𝑏𝑆 and 𝑏𝑟𝑒𝑎𝑙 for each parameter set provides that the determined value for 𝐼 is suitable for the analysis of additional datasets where the same base material is used.Figure 9 – Python based measurement versus manual measurement of widths.Calibration of the length measurementIn Figure 10, three frames from a video and their corresponding intensity profiles over the x-position in the FOV are shown. Based on the intensity profile at time 𝑡 = 𝑛2, the complete cycle of the exposure of a single track can be inferred. Material that has not been exposed by the laser exhibits an intensity value of I = 0 (𝑥 ≥ 1,000 µ𝑚). Material closer to the laser is preheated, leading to increasing radiation emission (890 µ𝑚 ≤ 𝑥 ≤ 1,000 µ𝑚).

Material closer to the laser is preheated, leading to increasing radiation emission (890 µ𝑚 ≤ 𝑥 ≤ 1,000 µ𝑚). Once the emitted radiation exceeds the sensor's capacity at a certain temperature, an intensity value of 255 is recorded. During the laser exposure and a certain trailing period, a plateau is formed with 𝐼 = 255 (600 µ𝑚 ≤ 𝑥 ≤ 890 µ𝑚). As the material cools down, the detected 𝐼 decreases, where the decrease can be approximated exponentially (0 µ𝑚 ≤ 𝑥 ≤ 600 µ𝑚).In Figure 10, 𝐼𝑚𝑒𝑙 is also represented as a red horizontal line. The value of 𝐼𝑚𝑒𝑙 = 149, which was determined iteratively in advance, is also used for evaluating the melt pool length. The figure illustrates how the melt pool length can be determined ∆𝑥𝑚𝑒𝑙 = 𝑥2 − 𝑥1 = 277.5 µm. Also,1127a new intensity value 𝐼𝑠𝑜𝑙 is introduced in Figure 10 as a blue horizontal line and describes the intensity at which the material solidifies after the liquid phase.

Also,1127a new intensity value 𝐼𝑠𝑜𝑙 is introduced in Figure 10 as a blue horizontal line and describes the intensity at which the material solidifies after the liquid phase. The need to use this value can be derived from Figure 11.2=500 µ1=0 µ 3=1000 µ1=0 µ2=500 µ 3=1000 µFigure 10 – Three intensity profiles, evaluated along the scan vector and used intensities to determine the melt pool length.It shows the results of the determined 𝑙𝑠_𝑚𝑒𝑙, 𝑙𝑠_𝑠𝑜𝑙 and 𝑙𝑟𝑒𝑎𝑙. Also depicted are three consecutive images taken during the exposure of a single track while illumination was activated which were used for the manual determination of 𝑙𝑟𝑒𝑎𝑙 values. Distinctive spots on the melt pool were analyzed and evaluated from the point when they no longer showed any change. In this case, a point with significantly increased intensity due to reflection of the illumination intensity was used as a distinctive spot. At 𝑡 = 0 µs, this point is not visible but appears at 𝑡 = 25 µs.

At 𝑡 = 0 µs, this point is not visible but appears at 𝑡 = 25 µs. Since in the next1128frame (𝑡 = 50 µs) and all the following ones no change of this point can be detected, it can be concluded, that the material is solidified. The mean value of 𝑙𝑟𝑒𝑎𝑙 = 533.12 µm ± 28.02 µm was determined with more than 15 measurements in different videos. When selecting the videos or areas to be measured, care was taken to ensure that a concise change in the melt pool geometry could be investigated in each case, as shown in Figure 10. This prevents subjective perceptions from having a major influence on the results. With the Python script and 𝐼𝑚𝑒𝑙 = 149, which was determined for the melt pool width, an 𝑙𝑆_𝑚𝑒𝑙 = 277.51 µm ± 40.21 µm was calculated. The high deviation between the manually and Python-based determined melt pool lengths of 255.61 µm or 52 % suggests that 𝐼𝑚𝑒𝑙 = 149 cannot be used for the detection of 𝑏𝑆 and 𝑙𝑆.

The high deviation between the manually and Python-based determined melt pool lengths of 255.61 µm or 52 % suggests that 𝐼𝑚𝑒𝑙 = 149 cannot be used for the detection of 𝑏𝑆 and 𝑙𝑆. Therefore, another intensity, referred to as 𝐼𝑠𝑜𝑙, was introduced, as shown in Figure 10. The length of the melt pool is subsequently calculated using the formula ∆𝑥𝑠𝑜𝑙 = 𝑥2_𝑚𝑒𝑙 − 𝑥1_𝑠𝑜𝑙. Thereby, the 𝐼 where material melt is chosen as before, just the intensity where material changes from liquid to solid state is corrected. After incorporating these conditions into the Python script and applying 𝐼𝑚𝑒𝑙 = 149 and 𝐼𝑠𝑜𝑙 = 24, the value of the lengths can be determined as 𝑙𝑆_𝑠𝑜𝑙 = 540.37 µm ± 36.72 µm.

After incorporating these conditions into the Python script and applying 𝐼𝑚𝑒𝑙 = 149 and 𝐼𝑠𝑜𝑙 = 24, the value of the lengths can be determined as 𝑙𝑆_𝑠𝑜𝑙 = 540.37 µm ± 36.72 µm. The deviation between 𝐿𝑠_𝑠𝑜𝑙 and 𝑙𝑟𝑒𝑎𝑙 is lowered with this approach to 7.25 µm or 1.4 %.=25 µ=50 µ==25 µFigure 11 – Results of manual and Python based determination of melt pool lengths with example images for the manual analyzation.The application of the methodology that the melt intensity is not equal to the solidification intensity may initially seem unusual. However, the varying intensities can be caused by the solidification range of an alloy. Alloys melt (liquidus) and solidify (solidus) at different temperatures, with the liquidus temperature being higher than the solidus temperature. Furthermore, mechanisms can arise that lead to an undercooling of the melt and thus increase the temperature respectively intensity difference [33, 34].

Furthermore, mechanisms can arise that lead to an undercooling of the melt and thus increase the temperature respectively intensity difference [33, 34]. Another factor contributing to the unequal intensity values could be the influence of the systems laser on the detected intensities, despite the1129high radiation filtration, compared to Figure 5. Interferences of melt intensities are also possible in the vicinity of the region where the laser exposes the material, as the irregular geometries of the melt pool promote such occurrences. A Validation of this method as described for the width is still pending at this time. However, the presented method offers a good approach for automated length determination.Cooling durationBesides evaluating the melt pool lengths, the cooling time Δ𝜏𝑥,𝑦 can also be calculated for each dataset with Δ𝜏𝑥,𝑦 = 𝜏𝐼_𝑠𝑜𝑙 − 𝜏𝐼_𝑚𝑒𝑙 . The result of such an evaluation, where parameter set 1 was used, is shown in Figure 12.

The result of such an evaluation, where parameter set 1 was used, is shown in Figure 12. With this figure, the cooling behavior of the intensities dependent on the position in x- and y-direction can be observed. Furthermore, it is possible to give information about the local intensity change and therefore relative temperature in the melt pool. Overall, the solidification of the material starts at the outer edge of the melt pool, where Δ𝜏𝑥,𝑦 is the smallest. As the pixel gets closer to the scan vector, the value of Δ𝜏𝑥,𝑦 increases, indicating a slower solidification. The white pixels around the data are those where the intensity 𝐼𝑚𝑒𝑙 was not exceeded during the examinations, so there is no data there.

The white pixels around the data are those where the intensity 𝐼𝑚𝑒𝑙 was not exceeded during the examinations, so there is no data there. Overall, a nearly symmetrical cooling can be assumed along the entire scan vector.,Figure 12 – Cooling duration 𝜟𝝉𝒙,𝒚 along a scan vector, exposure and recorded using parameter set 1.Defect analysisTo assess the system's ability to detect areas with anomalies, experiments were conducted where defects were intentionally induced. One specific defect type that is intentionally created is known as "balling." For this purpose, parameter set 2, compare section “Methods” is used. The evaluation of a dataset is shown in Figure 13. The number of frames Δ𝜏𝑥,𝑦 required between the melting of a pixel and its subsequent solidification or the first occurrence of exceeding 𝐼𝑚𝑒𝑙, followed by falling below 𝐼𝑠𝑜𝑙, is represented using a color scale. White pixels indicate that 𝐼𝑚𝑒𝑙 was not reached.

White pixels indicate that 𝐼𝑚𝑒𝑙 was not reached. From the figure, it is evident that the cooling durations in two areas are significantly longer (Δ𝜏𝑥,𝑦 >= 80 frames) compared to other areas (Δ𝜏𝑥,𝑦 ~ 40 - 60). Higher ∆𝜏 values typically indicate regions with increased melt volume. Between these two areas, marked with red circles, there is a region where the ∆𝜏 values are lower. In this region, it can be assumed that the melt volume is lower. Such variations along a melt pool can be an indicator of melt pool disruptions and thus the occurrence of the balling effect.1130,Figure 13 –Cooling duration Δ𝜏𝑥,𝑦 in frames of a dataset, gained by single track exposure with 𝑷𝑳 = 𝟑𝟐𝟎 𝑾 respectively parameter set 2.To investigate the hypothesis that the balling effect occurs at this location, the analyzed dataset is additionally evaluated manually. Selected frames from this dataset are shown in Figure 14.

Selected frames from this dataset are shown in Figure 14. To better visualize the observed effects, they have been processed in the Photron Fastcam Viewer software using the HDR tool at level 3, which corresponds to brightening the video. The exposure level is depicted. The laser enters the field of view (FOV) from the left side and exposes the material, which subsequently melts. The melt has a higher intensity than the surrounding powder. After the laser is no longer within the FOV, melted particles can be observed being ejected through the FOV (𝑡 = 2,250 µs). Subsequently, the intensity of the pixels decreases. At time 𝑡 = 3,600 µs, it becomes apparent that powder particles adhere to the melt, emitting significantly lower intensity. A first interruption of the intensity profile is noticeable. At 𝑡 = 4,850 µs, an elevated intensity is still visible in the left portion of the image.

A first interruption of the intensity profile is noticeable. At 𝑡 = 4,850 µs, an elevated intensity is still visible in the left portion of the image. Overall, with the analysis of these images the suggestion that the balling effect occurred within the scan track could be confirmed. Furthermore, it could be seen, that the powder particles, which are connected to the melt track at 𝑡 = 2,250 µs influence the determination of ∆𝜏, compare Figure 13 (left area where light blue pixels occur in the purple region) and Figure 14. This leads to the conclusion that even small influences in the melt pool region could be determined.Figure 14 - Emergence of the Balling Phenomenon on a Scan Vector1131Conclusion and OutlookIn this study, a method was developed to evaluate the melt pool width, melt pool length, and cooling duration of individual pixels recorded while generating melt tracks within the PBF- LB/M process.

To accomplish this, the objective of the high-speed camera system was brought into the build chamber of an SLM 280 Hl machine and hermetically sealed, allowing for the capture of videos with high spatial resolution (1.44 µm x 1.44 µm per pixel) and high temporal resolution (40,000 FPS). The analysis of the acquired datasets was performed using a custom-developed Python script, which involved the manual determination of the melt and solidification intensities 𝐼𝑚𝑒𝑙 through correlation. The melt intensity was determined based on the melt pool width and validated using the length parameter. The results demonstrate that the introduction and application of an 𝐼𝑚𝑒𝑙 and 𝐼𝑠𝑜𝑙 parameter in the Python script are effective in determining geometric properties.

The results demonstrate that the introduction and application of an 𝐼𝑚𝑒𝑙 and 𝐼𝑠𝑜𝑙 parameter in the Python script are effective in determining geometric properties. The melt pool width was determined with a deviation of 1.2 % from the manually measured width, and the melt pool length was determined with a deviation of 1.4% using the Python script and standard PBF-LB/M process parameters.Furthermore, it is possible to analyze cooling durations, which means the number of frames respectively time that a pixel intensity is above the defined 𝐼𝑚𝑒𝑙. With this, it is possible to determine irregularities within the melt track, for example, if the cooling duration of a region is lower/higher than expected. With this approach, it could be shown that the occurrence of balling could be detected on single tracks.Since the size of the datasets is quite high and the analysis of the data must be done in a sequential process, it cannot be used for real-time monitoring of the process yet.

On the other hand, there are quite a lot of potential use cases for this method. In the field of fundamental research, this method might be useful to gain more insight into the process behavior during the melting and solidification of novel materials. Here, materials, where it is needed to gain more knowledge about the dynamics of the melt track or local temperature changes, can be observed in future work. For example, materials such as metallic glasses or diamond-reinforced metal matrix composites can be suitable materials to be observed, since the local cooling durations have a huge impact on crystallization in bulk metallic glasses or the graphitization of the diamonds embedded in the metal matrix [35–38]. With improvements of the script, it could also be possible to track the movement of diamonds within the melt track. In addition, it is useful to correlate the intensities detected by the HSC system with real temperatures as Lane et al. [39] has performed.

In addition, it is useful to correlate the intensities detected by the HSC system with real temperatures as Lane et al. [39] has performed. For this purpose, it would be useful to couple the system with a high-speed pyrometer within the analysis room.AcknowledgementsThe camera used in this work was funded by the Deutsche Forschungsgemeinschaft (DFG, grant number WI 2118/17-1). The PBF-LB/M machine used for this work was funded with support from the University of Duisburg-Essen and the Deutsche Forschungsgemeinschaft (DFG, grant number INST 20876/341-1).References [1] DIN EN ISO/ASTM, “DIN EN ISO/ASTM 52900 PBF-LB/M: Additive Fertigung –Grundlagen – Terminologie,”[2] W. Associates, Wohlers report 2022: 3D printing and additive manufacturing global state ofthe industry. Fort Collins (Colo.): Wohlers Associates, 2022.1132[3]I. Gibson, D. Rosen, and B. Stucker, Additive manufacturing technologies: 3D printing, rapid prototyping and direct digital manufacturing.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceHigh-Resolution Electrohydrodynamic Jet Printing of Molten PolycaprolactonePatrick M. Sammons, Kira BartonUniversity of Michigan, Ann Arbor, MI 48109 Mechanical Engineering DepartmentAbstractPolycaprolactone (PCL) is a biocompatible and biodegradable polymer that is commonly used in drug delivery systems, medical structures, and tissue engineering applications. Typical ad- ditive manufacturing methods of PCL structures for tissue engineering applications either require harsh organic solvents or are only capable of producing relatively large feature sizes, which are not compatible with some of the desired applications.

Electrohydrodynamic jet (e-jet) printing, an additive manufacturing process which uses an electric ﬁeld to induce jetting from a microcapillary nozzle, is an attractive method for producing PCL tissue engineering structures due to the achiev- able resolution and the ability to print highly viscous inks. In this work, experimental investigation into the ability to print pure, molten PCL using the e-jet process is carried out. A characterization of the process inputs that yield suitable printing regimes is presented. Demonstration of the achiev- able resolution with e-jet printing is presented in the form of printed, high-resolution structures.IntroductionThe application of biocompatible, biodegradable, and bioresorbable materials for drug de- livery [8], medical devices [9], and tissue engineering applications [10] is garnering increased attention.

A popular subclass of these materials are polymers [1] which include polylactide (PLA), polycaprolactone (PCL), polyethylene glycol (PEG), in additions to many others. In particular, PCL has found numerous applications in tissue engineering applications due to its relatively low tensile strength, high elongation at breakage, and processability.For tissue engineering applications, inkjet and electrospinning (e-spinning) of PCL is popu- lar. In inkjet, because it is solid at room temperature and its viscosity when molten, PCL is typically carried with a solvent in order to enable printing [13]. However, these solvents can be harmful or the resulting PCL structure after the solvent has been removed does not display desirable prop- erties [13]. Alternatively, e-spinning can be used to manufacture large structures comprised of small ﬁber elements using pure PCL [11]. The major disadvantage of the e-spinning process stems from the unpredictable behavior of the ﬁlament produced.

The major disadvantage of the e-spinning process stems from the unpredictable behavior of the ﬁlament produced. It is therefore difﬁcult to control the placement of these ﬁbers which in turn limits the minimum feature size that can be successfully created with e-spinning. Further, there are a number of tissue engineering applications where it is advantageous to precisely place material and to create ﬁne features. Therefore, alternative methods for processing of PCL should be considered.Electrohydrodynamic jet (e-jet) printing is an emerging micro-/nano-scale additive manu- facturing (AM) process that shows promise for a variety of applications [12] including the manu- facture of conformal, printed electronics and biomedical structures. In e-jet printing, a voltage is applied between an ink-ﬁlled emitter and an attractor (typically a charged conductive microcapil- lary nozzle and a grounded conductive substrate, respectively).

In e-jet printing, a voltage is applied between an ink-ﬁlled emitter and an attractor (typically a charged conductive microcapil- lary nozzle and a grounded conductive substrate, respectively). The electric ﬁeld formed by the voltage potential induces an electric stress in the ink and causes a bulk deformation. At a critical2302level of the electric ﬁeld, the electric stress overcomes restorative surface tension stresses in the ink and a discrete volume of ﬂuid is ejected from the emitter towards the attractor. By adjusting the magnitude and duration of the electric ﬁeld formed by the voltage potential, micro- and nano-sized droplets can be created on a variety of substrates. While the underlying physics of both the e-jet and e-spinning processes are similar, the two processes differ in several ways.

While the underlying physics of both the e-jet and e-spinning processes are similar, the two processes differ in several ways. First, e-jet printing is performed with the nozzle relatively close (∼3-10 times the nozzle diameter in e-jet printing ver- sus ∼100 times the nozzle diameter in e-spinning) to the substrate. This enables ﬁner control of the placement of material on the substrate. Secondly, e-jet printing is operated in a ”stop-start“ manner where the voltage is modulated throughout the course of a build to fabricate features. This is in contrast to e-spinning where the voltage is held constant and the ﬁlament produced in e-spinning is continuously collected, typically on a mandrel. Figure 1 provides a schematic comparing the two processes. In e-jet printing, the ejection of material is periodic in time either by modulating the voltage (“Pulse”) or by operating in a regime where the jetting is intermittent and caused by a persistent high voltage (“DC”).

E-spinning is a constant process in terms of both the applied voltage and the material ﬂow.SubstrateCollectorPulseDCVoltageTimeVoltageTimeE-Jet PrintingE-SpinningFigure 1: Schematic of the e-jet and e-spinning processes with characteristic voltage signals.The rest of the paper is structured as follows. First, a description of the experimental system used to e-jet print molten PCL is given, including properties of PCL. Then, characteristics of the observed printing behavior is discussed. Finally, a summary, conclusion, and a discussion of future work is given.Experimental System DescriptionIn this section, details of the physical experimental system are given. Then, material prop- erties of the PCL used here are given. Finally, details regarding the expected limiting behavior while printing molten PCL are given.A. E-Jet Printer DescriptionThe e-jet printer used in this work is a custom built system developed within the Barton Research Group.

E-Jet Printer DescriptionThe e-jet printer used in this work is a custom built system developed within the Barton Research Group. The e-jet printer is comprised of a motion system, a print head, and a set of auxiliary components. The motion system consists of two stacked Aerotech stages (ATS15020-23031-02B) to enable X-Y motion with a vertical Z-axis mounted adjacent to the X-Y stage to enable vertical positioning. The horizontal stages are each repeatable to ± 2 µmThe print head consists of an in-house machined aluminum body into which a heating element and a syringe barrel can be inserted. The heating element is utilized with an external controller (Omron E5C2) and temperature feedback measured using a thermocouple attached to the aluminum body. To the syringe barrel, which contains the material to be printed, a custom fabricated nozzle is attached.

To the syringe barrel, which contains the material to be printed, a custom fabricated nozzle is attached. The nozzle consists of tungsten ﬁlament 35 gauge blunt needle (WPI NF35BL) inserted into a shortened 15 gauge blunt stainless steel needle. The print head is mounted to the Z-axis.There are three auxiliary systems integrated into the e-jet printer: a video camera, a high voltage ampliﬁer, and an air pressure regulator. The video camera (Lumenera Inﬁnity 2) is a color 2.0 Megapixel USB camera used to monitor the e-jet printing process. In line with the monitoring camera is a set of optics to provide magniﬁcation. The voltage ampliﬁer (Ultravolt/Advanced Energy HVA Series) produces up to 2 kV to the nozzle. Finally, the pressure regulator is capable of 1 PSI resolution at a minimum of 5 PSI. The staging, the pressure regulator, and the voltage ampliﬁer are controlled via a LabVIEW interface. Monitoring via the camera is enabled using software provided by the manufacturer.

The staging, the pressure regulator, and the voltage ampliﬁer are controlled via a LabVIEW interface. Monitoring via the camera is enabled using software provided by the manufacturer. Figure 2 shows a schematic of the e-jet printer with the above components and systems labeled.Figure 2: Schematic of the e-jet printer used in this work with relevant components and subsystems labeled (left) and a schematic of the custom nozzle fabricated for printing molten PCL (right).B. Polycaprolactone Material and Jetting PropertiesSelected relevant material properties of molten PCL are listed in Table 1 along with those of deionized (DI) water for comparison. Sigma-Aldrich polycaprolactone average weight number Mn = 45,000 is used for this work. The most noticeable difference between these two materials is viscosity.

Sigma-Aldrich polycaprolactone average weight number Mn = 45,000 is used for this work. The most noticeable difference between these two materials is viscosity. Molten PCL is approximately 1 × 106 more viscous than DI water, but possess similar electrical and surface tension properties.There are several unique challenges associated with attempting to print pure PCL. First, at room temperature, PCL is solid. Therefore, typical methods for printing PCL use one of several2304Table 1: Material properties of molten PCL and room temperature DI water.MaterialPCL ∼ 10 × 103 [7] Surface Tension Coefﬁcient, γ (N/m) ∼ 45 × 10−3 [4]Property Viscosity, µ (Pa-s)Density, ρ (kg/m3) Conductivity, σ (S/m) Relative Permittivity, (cid:15)r1145 2.6 × 10−5 [3] 3.5-4.5 [5]DI Water 1.0 × 10−3 72.8 × 10−3 998 0.1-5×10−5 ∼80solvents. Secondly, when molten, PCL is relatively viscous.

Secondly, when molten, PCL is relatively viscous. This is a signiﬁcant issue for systems such as inkjet and aerosol jet which possess an upper limit on ink viscosity well below that of molten PCL. Additionally, because molten PCL has a high viscosity, the time constant of the jetting behavior is governed by viscosity rather than surface tension.To understand the limiting behavior in the e-jet printing of molten PCL, consider the fol- lowing time scales. The capillary time scale of the process through a 150 µm nozzle, deﬁned as [6]τc =(cid:115)ρL3 γwhere τc is the capillary time scale (s), ρ is the ink density (kg/m3), L is the characteristic length (m) taken here as the nozzle inner diameter, and γ is the ink surface tension coefﬁcient (N/m), is τc = 2.93 × 10−4 s. Compare this value with that of DI water, τc = 2.12 × 10−4 s.

Compare this value with that of DI water, τc = 2.12 × 10−4 s. Alternatively, the viscous ﬂow time scale for a system with the same nozzle size, deﬁned as [2],τµ =ρL2 µwhere τµ is the viscous ﬂow time scale (s) and µ is the ink viscosity (Pa-s), is τµ = 2.58 × 10−8 s for molten PCL and τµ = 2.25 × 10−2 s for DI water. The viscous ﬂow time scale for molten PCL, where smaller time scales indicate greater importance in the ﬂow, dominates that of water. This is expected given the disparity in viscosity between the two ﬂuids.A signiﬁcant deviation of the behavior of molten PCL over typical inks in e-jet printing canbe observed by further examining the charge relaxation time,τe =(cid:15) σwhere τe is the charge relaxation time and an indicator of the importance of the charge density movement in the ﬂuid (s), σ is the ﬂuid conductivity (S/m), and (cid:15) is the ﬂuid permittivity (F/m).

For molten PCL, the charge relaxation time scale is τe = 1.02 × 10−6 s while that of DI water is τe = 3.54 × 10−5 s. In context, it can be seen that the viscous ﬂow scale signiﬁcantly dominates the other two time scales for PCL while the charge relaxation time scale dominates for DI water. However, there exists more parity in the time scales for DI water than for PCL.From the above analysis it can be expected that the jetting behavior and ﬂow of molten PCL will be limited by viscous ﬂow and that the jetting will happen over a longer time span than inks with similar material properties but lower viscosity.2305(1)(2)(3)E-Jet Printing CharacterizationIn order to characterize the jetting ability of molten PCL in the experimental setup de- scribed above, a series of tests are performed using pulse printing (see Figure 1). During the ex- ploration of this regime, evidence of a regime where DC printing results in intermittent jetting was discovered.

Pulse PrintingIn this section, details regarding the size of features produced using a select range of volt- ages and pulse widths will be presented. Table 3 gives the parameter levels tested here. The parameters presented in Table 3 were initially determined by exploring the voltage-pulse width space and noting the boundaries where the ﬂuid forms a distinct cone and successfully deposited material on the substrate.Table 3: Experimental levels used for characterizing PCL pulse printingParameter High Voltage (V) Pulse Width (ms) {1500,2000,2500}High 1000Level Mid 850 {3000,2500,2000}Low 750 {3500,4000,4500}Table 4 provides a summary of the Feret diameter, i.e., the largest dimension across the structure, statistics for each of the experimental conditions. Figures 3 through 4 give histograms and normal distribution ﬁts for selected data sets.

Figures 3 through 4 give histograms and normal distribution ﬁts for selected data sets. Examining the results in Table 4 and the selected results in the provided ﬁgures, it can be observed that the droplet diameter is approximately con- stant across the tested voltages and pulse widths. Further, while the mean values of diameter are all extremely close, the standard deviation for all of the tested combinations, with the exception of the ﬁnal 750 V trial, are less than 20% of the mean. This indicates very repeatable behavior.Figure 5 summarizes the trends of the measured Feret diameter with respect to pulse width and voltage. As stated above, the printing behavior appears insensitive to changes in pulse width and voltage.

As stated above, the printing behavior appears insensitive to changes in pulse width and voltage. While this is initially a surprising result, in the context of the time scales presented above, it is an expected result as the PCL behavior is dominated by the viscous ﬂow time scale rather than charge relaxation which is typically the case in e-jet printing.2306Table 4: Summary of pulse printing results. Experimental Condition Mean (µm) Std. Dev.

DC PrintingWhile exploring the voltage-pulse width parameter space, intermittent jetting was observed while voltage was high. This is a previously unreported behavior for inks as viscous as PCL.2307m)m)14161820222426283032Feret Diamter (60070080090010001100Voltage (V)14161820222426283032Feret Diamter (10002000300040005000Pulse Width (ms)Figure 5: Feret diameter as a function of voltage (left) and pulse width (right) for all tests.Further, it provides a regime in which features smaller than those observed above could be made without reducing the nozzle size as is tpyical in e-jet printing to achieve higher resolution. Figure 6 shows still images taken from a video demonstrating the intermittent jetting behavior. During the 2000 ms pulse width, seven jetting events were observed. In frames labeled t =743, 908, 1238, 1321, and 1486 through 1651 ms, i.e., labeled “Jet”, a sharp meniscus can be observed.

During the 2000 ms pulse width, seven jetting events were observed. In frames labeled t =743, 908, 1238, 1321, and 1486 through 1651 ms, i.e., labeled “Jet”, a sharp meniscus can be observed. While not entirely clear in the still images, primarily due to the low frame rate of the monitoring camera, a jet of material issues from the meniscus during these frames. In between the frames depicting jetting, the meniscus retracts before extending again into a sharp cone.NozzleReflectiont = 1981 msJetJetJett = 580 msPCLt = 0 ms135 μmt = 1321 mst = 908 msJetJetJett = 660 mst = 743 mst = 1403 mst = 826 msJetChargingDischargingt = 1734 mst = 1569 mst = 1486 mst = 1651 mst = 991 mst = 1238 msFigure 6: Still images taken from the system monitoring camera showing evidence of an intermittent jetting mode for pure PCL.2308ConclusionPolycaprolactone (PCL) is an important polymer for tissue engineering applications.

The typical methods for fabrication of PCL structures is through either inkjet printing or e-spinning. While both of these processes have beneﬁts for speciﬁc applications of PCL structures, there exists a gap in fabrication methods for small patterns of ﬁne features. E-jet printing, a process governed by the same physics as e-spinning, is a promising tool for ﬁlling this gap. Here, e-jet printing is used to demonstrate the ability to print high-resolution features using molten PCL.In the drop-on-demand or pulse printing regime, where a pulsed voltage is applied to the printing nozzle, voltages and pulse widths ranged from 750 to 1000 V and 1500 to 4500 ms. These voltages are similar to those used in printing of other less viscous inks. However, the pulse widths are approximately two orders of magnitude greater than those typical for e-jet printing in the pulse printing regime.

However, the pulse widths are approximately two orders of magnitude greater than those typical for e-jet printing in the pulse printing regime. This property is a direct result of the limiting viscous ﬂow for molten PCL while other ink materials are limited by the charge relaxation time. While the processing window for molten PCL is similar to that of other inks typically e-jet printed, the typical relationships between voltage and pulse width and feature volume do not appear to hold. In fact, the process appears insensitive to the magnitude of voltage applied or the duration of the applied voltage within a range. Features of approximately 20 µm were produced for all of the experimental conditions used here. This further corroborates the viscosity limiting behavior of e-jet printed molten PCL.During initial testing, a regime was found in which intermittent jetting was observed. This is a previosuly unreported regime for PCL.

GH-SPEED CURTAIN RECOATING FOR STEREOLITHOGRAPHYM. Gilio, J.-P. Kruth, and P. VanherckDivision of Production Engineering, Machine design and Automation Katholieke Universiteit Leuven, Belgium 3001 HeverleeAbstractThe University of Leuven uses a liquid curtain recoating system for resin deposition in stereolithography. This system deposits new liquid layers of photo-polymer by means of a liquid curtain travelling over the build vat. Experiments have been carried out to increase the speed of the liquid curtain while depositing a layer. Speeds up to 1.2 m/s, and accelerations up to 1 g have been tested successfully, meaning that it is possible to coat high-quality layers of 75 µm thickness with this recoating technique. However, the curtain restores too slowly after acceleration. This paper discusses possible reasons and tries to formulate adequate solutions. Possible solutions consist in controlling small pressure differences in the curtain’s neighbourhood.

This paper discusses possible reasons and tries to formulate adequate solutions. Possible solutions consist in controlling small pressure differences in the curtain’s neighbourhood. A solution to this problem is necessary, as to make the travelling length of the curtain, and so the machine length acceptable with respect to the dimensions of the build vat.IntroductionThe process of curtain coating consists of pumping a coating liquid from a reservoir into a precision extrusion head, which has a narrow slot along its lower face. The liquid is extruded out of the slot, and in ﬂowing downwards under the inﬂuence of gravity it forms a liquid sheet, also called a liquid curtain. The liquid curtain then hits a substrate. This substrate moves relative to the curtain, and by so doing a liquid ﬁlm is coated onto it. Liquid sheets are often employed to deposit uniform liquid layers on a moving substrate as, for example, in the coating of photographic ﬁlm.

Liquid sheets are often employed to deposit uniform liquid layers on a moving substrate as, for example, in the coating of photographic ﬁlm. For more applications, the reader should consult Finnicum [4] and the references of Brown [2]. Typically, the liquid curtain is long and thin, and can be up to several metres wide. Because of their technological and theoretical interest, the ﬂow in a liquid curtain has been extensively studied, addressing both the shape of the curtain [4, 5], and instability issues [8, 7, 2].In our application the extrusion head is displaced, rather than the substrate. The substrate, consisting of a build vat ﬁlled with liquid photo-polymer and a partially built stereolithography part, is not moving. In order to apply a new liquid layer in the stereolithography process, the extru- sion head together with the liquid curtain is moved over the build vat.

In order to apply a new liquid layer in the stereolithography process, the extru- sion head together with the liquid curtain is moved over the build vat. As several subsequent layers are coated during the process of fabricating a part with stereolithography, we refer to this coating technique as curtain recoating [9, 6]. Curtain recoating is one of the fastest coating techniques,46able to coat a new liquid layer for stereolithography in about two seconds. Considered the appre- ciable amount of layers a stereolithography product usually consists of, curtain recoating enables a substantial reduction of part build times.Basically, a build vat with moving platform is used as depicted in ﬁgure 1. Before coating a new layer of liquid photo-polymer, the platform is lowered over a distance equal to the layer thickness. Hereafter, the coating head, and thus the liquid curtain is moved over the build vat.

Hereafter, the coating head, and thus the liquid curtain is moved over the build vat. As a result, the liquid level in the build vat after each newly coated layer remains at constant height. A basin around the build vat collects the excess of resin and feeds it to a supply reservoir which, in turn, feeds a metering pump. In a sense, this application of curtain coating is unique with respect to all previously published studies, because of the coating head, rather than the substrate, is moving. The problem of seeking optimal operating conditions, necessary to coat layers of good quality is described in the following section.

The subsequent sections discuss the experiments that have been performed in order to gain a better understanding of the curtain’s dynamic behaviour, a necessary prerequisite for process improvement.build vat with movable platformcoating headliquid curtain(deceleration)(acceleration)direction of movementregion of constant speedacceleration lengthdeceleration lengthwind shieldgear pumpcollecting basinresinreservoirFigure 1: Curtain recoating for stereolithography: coating of a new liquid layerProblem descriptionThe thickness d of a coated liquid layer relates to the ﬂow rate q through the extrusionhead’s slot per unit slot width [m2/s], and the speed of the coating head U [m/s], according to:d =q U.In order to coat a layer of uniform thickness, q must be constant along the width of the curtain, and U must remain constant while the head crosses over the build vat.

There exists a minimum value for q for good operation, based on a stability issue for the liquid curtain. Below this minimal q the liquid curtain breaks and no intact liquid layers can be coated. Hence, in order to further decrease the thickness of a coated layer, U has to be increased. A typical value for q = 90 mm2/s for our coating head, meaning that in order to coat a layer of d = 100 µm, U must be 0.9 m/s; for d = 60 µm, U = 1.5 m/s.47(1)Actually, decreasing the layer thickness is not the primary goal of high-speed curtain re- coating. The deposition of a liquid layer of good quality is a more stringent problem to be solved. If the recoating speed U is equal to the ﬂuid velocity V in the curtain at a point where it impinges onto the substrate, then the thickness d of the coated layer will be equal to the thickness h of the curtain at its lower end. When U is smaller than V , d will be larger than h, and a kind of ‘heel’ will form where the curtain hits the substrate [1, 3].

When U is smaller than V , d will be larger than h, and a kind of ‘heel’ will form where the curtain hits the substrate [1, 3]. Figure 2 depicts this phenomenon. Recirculating ﬂows will develop in this heel, thereby entrapping air in tiny bubbles. These bubbles will, eventu- ally, be trapped in the liquid layer. At present, the curtain in our machine is 125 mm long, and the velocity of the ﬂuid at the lower end of the curtain is about 1.4 m/s [2]. The maximum achievable speed U in our machine is 0.7 m/s, meaning that the thickness of the coated layer d is about two times the curtain thickness h at its lower end.

The maximum achievable speed U in our machine is 0.7 m/s, meaning that the thickness of the coated layer d is about two times the curtain thickness h at its lower end. Thus, our primary goal in this research is seeking better operating conditions than the ones presently used.U<V, formation of a heelnormal operation: U/V = 11heelcoating speed UUfluid velocity VFigure 2: Formation of a heel at the lower end of a liquid curtain when the coating speed is substantially smaller than the ﬂuid velocity in the curtainIn order to coat a liquid layer, the coating head has to be accelerated from rest to the desired coating speed U . Throughout this paper the distance travelled by the head during acceleration will be called the acceleration length (ﬁg. 1). A section of constant speed U follows next. In the be- ginning of this constant speed section, the curtain has to reassess from acceleration. Subsequently, a liquid layer is coated.

1). A section of constant speed U follows next. In the be- ginning of this constant speed section, the curtain has to reassess from acceleration. Subsequently, a liquid layer is coated. Finally the coating head decelerates in order to be at rest at the end of its stroke. When coating the subsequent layer, the head starts at this latter position and moves towards the former rest position, thus necessitating a symmetrical machine.Some problems that are encountered in high-speed curtain recoating are now discussed. Firstly, the machine dimensions have to remain within reasonable proportions. It would be un- acceptable to have a machine of more than 3 m long with a build vat of only 250 mm diameter. The acceleration length should thus be limited. When increasing the coating head’s speed U , the peak acceleration should be increased, rather than the acceleration length. Secondly, inertia forces during acceleration will bend the liquid curtain backwards.

Secondly, inertia forces during acceleration will bend the liquid curtain backwards. Vertical rods at both lateral ends of the curtain are needed to maintain a straight curtain, as surface tension would otherwise contract it. The curtain is pinned at these rods, also called ‘edge guides’ (ﬁg. 3). Thus, when bending back- wards the curtain will curve also in a horizontal plane, remaining pinned at the edge guides and assuming a maximal bending in its middle. This situation is depicted in ﬁgure 4. After reaching the desired coating speed U , the curtain slowly returns to its original, planar shape. This planar shape has to be reached before beginning to coat the layer, otherwise the contact line between curtain and48substrate will be curved. Since, in this case, the curtain is still in a transient state, its velocity is not equal to the constant speed U , and the thickness of the coated layer will not be uniform.

Since, in this case, the curtain is still in a transient state, its velocity is not equal to the constant speed U , and the thickness of the coated layer will not be uniform. A number of experiments have been performed in order to investigate the feasibility of using high-speed cur- tain recoating for stereolithography. The following section describes the set-up that has been used for the experiments.Experimental set-upA number of experiments have been performed in order to investigate the possibility of high-speed curtain recoating. In the context of curtain recoating for stereolithography, high-speed means coating speeds U higher than 1 m/s. The coating head was mounted on a linear drive, having a total stroke of 900 mm. This linear drive was chosen because of its ease of programming different acceleration proﬁles, and because of the high accelerations the drive can achieve. Degraded Somos 7100tm resin with a viscosity of 1.1 Pa.s @ 34◦C was used for the experiments.

Degraded Somos 7100tm resin with a viscosity of 1.1 Pa.s @ 34◦C was used for the experiments. The curtain was 400 mm wide and 125 mm long. Figure 3 shows the set-up. A Sony digital video camera was mounted next to the coating head, moving together with the head and slide. Video frames were recorded at a rate of 50 Hz during motion.video cameraliquid curtainedge guidescoating headfeed pipecollecting basinlinear driveperforated plateFigure 3: Experimental set-up for the investigation of the curtain’s behaviour during acceleration to high speeds of the coating head.Different acceleration proﬁles were tested, among these a 2nd order polynomial, having a discontinuous jerk proﬁle at the beginning and end of the acceleration section, and a 4th order polynomial, having a continuous jerk proﬁle throughout the entire trajectory of the drive. For a given acceleration length, the 2nd order proﬁle gives the lowest peak acceleration.

For a given acceleration length, the 2nd order proﬁle gives the lowest peak acceleration. For a given peak acceleration, the 2nd order proﬁle gives the shortest acceleration length. No substantial difference has been observed in the curtain’s response to both mentioned proﬁles. This observation is also supported by the results of a simulation of the curtain’s behaviour during acceleration of the coating head. As a liquid curtain possesses a very low stiffness, a shield is mounted at both sides of the coating head to protect the curtain from aerodynamic forces during motion. Figure 5 and 4 show cross-sections of a backwards bent curtain. The following section covers the performed experiments in more detail, and discusses the results.49direction of movementshieldbent curtainposition of the curtain in steady−state and of the slotedge guideshielddirection of movementcoating headshieldmax.

deflectionliquidcurtainslotFigure 4: Horizontal cross-section of a liquid curtain bending backwards during motion. The shields are protecting the curtain from aerody- namic inﬂuences.Figure 5: Vertical cross-section of a falling liquid curtain bending backwards during mo- tion.Results and discussionA ﬁrst issue in the experiments was to observe up to which maximum peak accelerations the curtain could resist without breaking. A higher peak acceleration shortens the acceleration length needed to achieve a desired speed U , thus making the machine shorter. The coating head and curtain were subjected to a peak acceleration of 1g (9.8 m/s2), both with a 2nd and 4th order polynomial acceleration proﬁle to reach a speed U of 0.9 m/s only after an acceleration length of respectively 62 mm and 77 mm for the 2nd and 4th order polynomial proﬁle. During acceleration, maximum backwards bending of the curtain, as deﬁned by ﬁgure 5, was 42 mm. No breakage of the curtain was observed.

During acceleration, maximum backwards bending of the curtain, as deﬁned by ﬁgure 5, was 42 mm. No breakage of the curtain was observed. These observations show that high accelerations are no limiting factor for high-speed curtain recoating. Secondly, the feasibility of letting the curtain move at speeds higher than 1 m/s was investigated. Due to the limited stroke of the linear drive, the maximal speed tested was U = 1.2 m/s. To reach this speed, a peak acceleration of 9.0 m/s2 was used and the acceleration length was respectively 120 mm and 150 mm for the 2nd and 4th order polynomial proﬁle. No breakage of the curtain was observed, and the curtain’s maximum backwards bending was 52 mm with respect to the slot’s position.However, in restoring towards its ﬂat shape after acceleration to a speed U , the curtain takes a much longer time than would be expected.

As the curtain must return to its ﬂat shape before actual coating can start, the observations show that extra length has to be added to the travelling length of the coating head, in order to give the curtain sufﬁcient time to reassess. It is the purpose of this paper to investigate the reasons of this unwanted slow reassessment of the curtain, and to formulate a possible solution. The behaviour of the curtain during acceleration of the coating head has been simulated. A two-dimensional model of the curtain has been used for simulation, thus assuming a curtain of inﬁnite width. As the curtain in reality, is pinned laterally to the edge guides, it reaches its maximal bending in the middle. This pinning of the curtain at its lateral ends will undoubtedly inﬂuence the amount of bending in the middle, as the curtain is not free to deﬂect50as a whole. Therefore, the simulation will always yield a higher deﬂection than experimentally observed.

Therefore, the simulation will always yield a higher deﬂection than experimentally observed. It is not a purpose of this paper to describe the simulation in detail. What does matter in this stage of analysis, is that the model is able to qualitatively predict the experimentally observed behaviour of the curtain.The curtain’s shape is governed by the equations of Navier-Stokes for viscous ﬂow, suitably transformed to be used for liquid sheets. External forces acting on the curtain are: gravitation, inertial forces due to the acceleration, and aerodynamic forces induced by the air between the curtain and the two shields. The inﬂuence of viscosity and surface tension was also modelled. The aerodynamic forces, that act as a pressure difference between the curtain’s front and back side, are most subject to uncertainty.

The inﬂuence of viscosity and surface tension was also modelled. The aerodynamic forces, that act as a pressure difference between the curtain’s front and back side, are most subject to uncertainty. The reason for this uncertainty is due to the fact that, although the shields indeed protect the moving liquid curtain from still air, an amount of air still enters the space between curtain and shield through a gap between the lower edge of the front shield and the substrate. The shields may not contact the liquid substrate. Therefore the shields are mounted such that a gap of 2 mm is left between them and the substrate. When the curtain is moving, air enters through this gap into the space between curtain and shield. The liquid curtain will subsequently have to deﬂect the inﬂowing air toward its lateral ends, where the air is free to escape.

The liquid curtain will subsequently have to deﬂect the inﬂowing air toward its lateral ends, where the air is free to escape. Since in this case the curtain is moving at a speed U , while the air can be considered to be at rest, the aerodynamic forces have been modelled as a stagnation pressure, acting as a pressure difference between the curtain’s front and back side. The modelled stagnation pressure relates to the coating speed U according to:ps =ρairU 2 2with ρair the density of air. This pressure is very low (≈ 1 Pa when U = 1.2 m/s), but its inﬂuence on the curtain dynamics is nevertheless tremendous, because the bending stiffness of the curtain is also very low. According to our model, and considering a steady state, a constant pressure difference of 1 Pa between the curtain’s front and back side causes the curtain to deﬂect by 70 mm at its lower end, measured from the slot’s position.

The reader should bear in mind that the length of the curtain in this case is only 125 mm.With this in mind, and in order to qualitatively predict the experimentally observed be- haviour, the reader should notice from ﬁgures 6 and 7 that during the motion at constant speed U , the curtain shows a relaxation towards its planar shape. The relaxation obviously suggests that the pressure difference across the curtain slowly vanishes. This means that the stagnation pressure in our model should be subject too to a relaxation, which was implemented as an exponential decay of the stagnation pressure. As the dynamics of the experimentally observed relaxation are not clear yet, the relaxation constant of the modelled exponential decay of the stagnation pressure is found such that a best ﬁt between the experimental data and the model’s results is achieved.

A relaxation constant τ ≈ 0.2 sec is found to agree best with all data, as can be seen in ﬁgures 6 and 7.The introduction of a stagnation pressure is open for discussion; what is important though, is that the experimental data shows that the curtain ﬁrst deﬂects, due to inertial and aerodynamic forces, the latter ones increasing with increasing speed U . The deﬂection shows afterwards a kind of relaxation. This relaxation is mainly caused by a pressure in front of the curtain, that has built up during acceleration and is now disappearing slowly. A ﬁrst possible explanation for the dynamics51(2)−0.0150.30.51max curtain deflection [m]accelerationvelocityacc. [x10 m/s2] ,vel.

A ﬁrst possible explanation for the dynamics51(2)−0.0150.30.51max curtain deflection [m]accelerationvelocityacc. [x10 m/s2] ,vel. [m/s]time [sec]Maximum curtain deflection [CS09−230305]0.80.10.5020.20.70.60.10010.40.9−0.030.50.20.90.800.4simulationexperiment0.71.510.3−0.060.6−0.045Figure 6: Plot of the curtain’s deﬂection measured at its lower end with respect to the slot’s po- sition. Both experimental data and simulation results are plotted; values refer to left y-axis. The applied acceleration (2nd order polynomial proﬁle, amax = 2.7 m/s2) is plotted, together with the speed of the coating head Umax = 0.9 m/s (right y-axis).acc. [x10 m/s2] ,vel.

The applied acceleration (2nd order polynomial proﬁle, amax = 2.7 m/s2) is plotted, together with the speed of the coating head Umax = 0.9 m/s (right y-axis).acc. [x10 m/s2] ,vel. [m/s]time [sec]Maximum curtain deflection [EB12−280308]0.9accelerationvelocity0.20−0.10.70−0.0250.40.2simulationexperimentmax curtain deflection [m]0.60.30.3simulationexperiment0.60−0.0500.50.50.90.510.1−0.0750.120.81.50.7110.80.4Figure 7: Plot of the curtain’s deﬂection measured at its lower end with respect to the slot’s po- sition. Both experimental data and simulation results are plotted; values refer to left y-axis. The applied acceleration (4th order polynomial proﬁle, amax = 9.0 m/s2) is plotted, together with the speed of the coating head Umax = 1.2 m/s (right y-axis).52of the relaxation is a rearrangement of the air ﬂow in the space between the curtain and the air shield.

Once a constant speed U has been achieved, a rearrangement of the air ﬂow could drop the pressure in the direct vicinity of the curtain’s front. A second possible explanation is a pressure building up behind the curtain, thereby balancing the pressure in front of the curtain. The dynamics of this pressure build-up could reasonably be slower than the ones accounting for the curtain’s front pressure. The pressure differences occurring in this problem are very low (magnitude 1 Pa). Since pressure differences are the driving force behind ﬂows, the ﬂow rate of air is minimal too, thereby serving as a possible explanation for the slow relaxation. Moreover, the compressibility of air certainly affects the curtain’s dynamics, because the speed of propagation of a pressure disturbance is ﬁnite in a compressible medium.

Moreover, the compressibility of air certainly affects the curtain’s dynamics, because the speed of propagation of a pressure disturbance is ﬁnite in a compressible medium. The relaxation constant τ , that has been introduced in our simulation is thus a generic parameter, accounting for all physical effects that inﬂuence the observed relaxation.A validation of these assumptions is difﬁcult, since the pressure measurements are difﬁ- cult to perform because of their small magnitude and because of the occurring accelerations that will inﬂuence the measurements. Another approach consists in visualising the streamlines of air around the curtain in a wind tunnel. Such a test would be incomplete though, due to the inabili- ty to simulate inertial forces acting on the curtain. Actually, the assumptions could be tested by experiments aiming at inﬂuencing the relaxation.

Actually, the assumptions could be tested by experiments aiming at inﬂuencing the relaxation. If the relaxation is dependent upon a presumed pressure difference between the curtain’s middle and its lateral ends, the speed of relaxation could be increased by increasing the mentioned pressure difference. A second solution could be accom- plished by increasing the pressure at the back of the curtain, thereby balancing in a much faster way the stagnation pressure that has built up in front of the curtain. Anyway, it is clear that there are methods to dynamically control the curtain’s deﬂection during acceleration. These ideas have not been implemented yet, but will be a research topic in the future.

Anyway, it is clear that there are methods to dynamically control the curtain’s deﬂection during acceleration. These ideas have not been implemented yet, but will be a research topic in the future. They are needed for im- plementing high-speed curtain recoating for stereolithography while maintaining the length of the coating head’s drive, and thus the machine length within reasonable proportions.ConclusionThis paper has discussed preliminary experiments for the implementation of high-speed curtain recoating for stereolithography. Curtain recoating is a fast technique to coat even very thin liquid layers. The terminology ‘high-speed’ is used in this application to refer to coating speeds higher than 1 m/s. Speeds up to 1.2 m/s and accelerations up to 1g have been tested successfully, without breakage of the curtain. A main problem to be solved is the curtain’s slow returning to its ﬂat shape after bending due to acceleration.

A main problem to be solved is the curtain’s slow returning to its ﬂat shape after bending due to acceleration. In this respect, the curtain’s dynamics have to be made faster. Possible solutions to this problem have been mentioned, all consisting in controlling small pressure differences in the curtain’s neighbourhood. A solution is necessary as to make the machine length acceptable with respect to the dimensions of the build vat.AcknowledgementsThe present research is ﬁnanced by the Institute for the Promotion of Innovation by Science and Technology in Flanders (IWT).53References[1] T.D. Blake, A. Clarke, and K.J. Ruschak. Hydrodynamic assist of dynamic wetting. AIChEJournal, 40(2):229–242, 1994.[2] D.R. Brown. A study of the behaviour of a thin sheet of moving liquid. J. Fluid Mechanics,10:297–305, 1961.[3] A. Clarke. Recirculating ﬂows in curtain coating. In 1st European Coating Symp. Med. Thin.Films, pages 32–41, 1995.[4] D.S. Finnicum, S.J.

Histogram - Based Algorithm for Semiautomated Three-Dimensional Craniofacial Modelingby Richard A. Levy, M.D.Department of RadiologyUniversity of Michigan HospitalsIntroduction: Volume averaging artifacts in medical imaging result from voxel occupancy bymore than one tissue type and, with anisotropic voxels, may be decreased bychanging the imaging plane orientation relative to the target tissue and/or bydecreasing slice thickness. 1In craniofacial CT imaging, volume averaging artifact becomes significant in areasof thin bone such as the orbital walls and auditory ossicles. These regions arecustomarily imaged using multiple scan planes and the thinnest slices possible toreduce such artifacts.

These regions arecustomarily imaged using multiple scan planes and the thinnest slices possible toreduce such artifacts. In three-dimensional craniofacial imaging, these sameparameters may be controlled to reduce partial volume averaging, but areas of bone"drop-out" (also called pseudoforamina) are commonly present secondary to aparadoxical inability to lower thresholds without including unwanted backgroundtissues.

At present, the optimal solution to this problem is achieved by manually(and often painstakingly) drawing a region of interest around tissues presumed tocontain volume averaged target density voxels and lowering thresholds to includethese voxels in the 3D reconstruction, one CT slice at a time.Recently, anatomic modeling technologies have demonstrated the feasibility ofassembling particulate hydroxyapatite (SYnthetic bone) into detailed craniofacialmodels of high anatomic accuracy, theoretically suitable for in vivo implantation(work in progress with the Department of Mechanical Engineering, University ofTexas at Austin.) These modeling systems, such as stereolithography and selectivelaser sintering, operate as do 3D imaging workstations, using thresholds to198Materialsand Methods:include/exclude pixels from CT data sets in the modeling process.

However, theuser interactive capabilities of such technologies may be limited such that manualtracing of regions of volume averaged thin bone is not possible. Drop-out artifactsin models so generated would be potentially larger than on corresponding 3Dimages where user input could reduce these artifacts. To circumvent this inability tomanually correct volume averaging artifacts on anatomic modeling systems, and torelieve the intensive oPerator input required to otherwise achieve this goal on 3Dimaging software, a histogram-based algorithm for semiautomated threedimensional craniofacial modeling was develoPed.CT data sets from 5 axially and coronally-scanned embryologiccraniofacial specimens (aged 18 to 33 weeks) were used to generate onedimensional histograms of entire imaging volumes. Histograms were constructedwith the y-axis representing frequency and the x-axis representing CT number inHounsfield units.

Histograms were constructedwith the y-axis representing frequency and the x-axis representing CT number inHounsfield units. Superimposition of histograms from corresponding orthogonalCT data sets from each case tested the hypothesis that voxel anisotropy producesdivergence of corresponding segments of the histograms over ranges of Hounsfielddensity in which partial volume averaging occurs.

These regions of volumeaveraging were predicted from prior phantom trials based upon 1) divergence of thehistograms and/or presence of an asymmetric "tail" in an otherwise symmetrictissue distribution 2) very small numbers of pixels compared to the largepopulations of bone, soft tissue/water and air and 3) locations intermediate inHounsfield density to the standard tissue distributions (Figures la and Ib).Thresholds based on these assumptions and on prior experimentation with this CTphantom applied to the histograms were compared with CT image-derived operatorselected thresholds.

Most bone density pixels were identified using a simpleformula obtained from prior phantom trials, T=O.16ID-DoI+Do, where199Results:T=threshold, D=mean of bone pixel distribution, and Do=mean of soft tissuedistribution.2 In cases in which volume averaging was predicted from thehistograms, the corresponding threshold ranges were expanded and contracted toassess the "uniqueness" of the predicted threshold ranges. All 3-dimensional CTimages were generated using a volumetric rendering algorithm. Pathologiccorrelation consisting of craniofacial specimen photographs was available in somecases.Only in 1/5 cases (the 33 week specimen) could bone, soft tissue/water and airpopulations all be identified from the histograms. In the remaining 4/5 cases, onlybone and air "peaks" were identified, with soft tissue/water voxels distributed morebroadly over less well-defined Hounsfield density ranges.

In the remaining 4/5 cases, onlybone and air "peaks" were identified, with soft tissue/water voxels distributed morebroadly over less well-defined Hounsfield density ranges. The ability to setthresholds was related to the ease with which the means of these tissue populationscould be identified.In 3/5 cases, inspection of the histograms permitted identification of discrete rangesof Hounsfield density over which partial volume averaging occurred. Asanticipated, inclusion of non-osseous volume averaged voxels in the 3Dreconstructions also occurred, principally at air-tissue interfaces such as the skinsurface and surface of mucous membranes (Figure 2). In all cases in which volumeaveraging was predicted from the histograms, expansion of the threshold rangesbeyond those predicted yielded inclusion of non-osseous tissue in the 3Dreconstructions.

In 2/3 of these cases, contraction of the threshold ranges causedloss of bone-density voxels from anatomic regions where bone was observed tofill-in using the predicted threshold ranges (Figure 2). In the remaining case,pathologic verification was not available.200Discussion: While volume averaging artifacts may be tolerated in 3D craniofacial imaging, theybecome more critical in craniofacial modeling, especially with the expectation thatbiosynthetic (e.g., hydroxyapatite) implants will promote osteoneogenesis.3,4Tolerance of "drop-out" artifacts in a craniofacial implant will depend on the sizeand location of such defects.The histogram-based thresholding algorithm presented in this essay proposes todiminish volume averaging artifact as well as the operator time involved in thegeneration of craniofacial implants from CT data.

At present, the rate limiting factoris the time required for data transfer from the CT console to the generation ofhistograms.Optimally, this process results in a single threshold on the order of 150 HU, as wellas a range of lower CT density, e.g., -100 to -400 HU (Figure 2). For the purposeof anatomic modeling, undesired volume averaged tissues such as the skin surfacecan be dissected from the implant.A potential pitfall to this method is its dependency upon the presence of welldefmed bone, water/soft tissue and air-density histogram populations to permit theinitial selection of thresholds. It is likely that the small specimen size used in thisexperiment contributed to the absence of defmable water/soft tissue peaks in mostcases, especially since the expected tissue distributions were achieved only with thelargest (33 week) specimen.

While the utilization of such small craniofacialspecimens has the advantage of eliminating (potential positionally-dependent) beamhardening artifacts, it is anticipated that further trials with this method will occurwith postnatal and adult subjects.Another potential pitfall awaiting evaluation in clinical trials of this algorithm arisesfrom the fact that, in the present experiment, the imaging volume encompassing201the entire craniofacial specimen was included in both the paired axial and coronaldata sets. In the clinical setting, this may not be possible, making identification ofvolume-averaged voxels more challenging.References1.Levy RA, Edwards WT, Meyer JR, Rosenbaum AB. Facial trauma and 3-D reconstructiveimaging: insufficiencies and correctives. AJNR 1992;13:885-892.2.Levy RA. Semiautomated three-dimensional craniofacial imaging and modeling.Investigative Radiology 1994;29(2): 150-155.3.Damien CJ, Parsons JR, Benedict JJ, Weisman DS.

Top~, Top i = operator-selected thresholds for 3Dreconstructions generated from axial and coronal data, respectively. Note gradualupslope ("tail") as well as slight divergence of the histograms in the intervals defined byTcalc and Top, with the histogram generated from coronal data (i) having the largernumber of pixels. The intervals defmed by Tcalc and Top were demonstrated tocontain volume averaged voxe1s in submillimeter structures in the phantom that "filledin" using the lower thresholds, Top. These structures were better visualized on 3Dimaging using the axial (~) data at a numerically larger Top, consistent with lessvolume averaging in the axial (~) 2D CT data set.202Figure 1bSuperimposed one dimensional histograms generated from axial and coronal 2D crdata sets of a 33 week craniofacial specimen with holoprosencephaly. Note "inversion"of the histograms between -100 and -400 HU, with the coronal data set having thegreater number of pixels/voxels.

Note "inversion"of the histograms between -100 and -400 HU, with the coronal data set having thegreater number of pixels/voxels. Volume averaged voxels would be expected to appearon 3D reconstruction using an additional threshold of -100 to -400 HU, with greater"fill_in" on the 3D reconstruction generated from coronal data.Figure 2aPhotograph of the 33 week holoprosencephalic specimen. Note intact frontal bones.Figures 2b,2c 3D reconstructions from axial (2b) and coronal (2c) data using operator-selectedthresholds of 183 and 186, respectively. Note better "flil-in" of frontal bone voxelsusing axial data.Figure 2d3D reconstruction from coronal data using the histogram-derived threshold of 130.Figure 2e3D reconstruction from coronal data using the combined thresholds of 130 and -100 to400.

Note "fill-in" of frontal bone voxels(*) as well as other volume averaged tissuessuch as the skin surface (arrow) and mucous membrane surface (double arrows).Figure 2fSame as in 2e but using axial data, illustrating inclusion of skin surface and mucousmembrane volume averaged voxels to a similar extent as with the coronal data set. Thisindicates that the region of histogram divergence between -100 to -400 HU in Figure 1bwhere the coronal data set contains a greater number of voxels corresponds to volumeaveraged frontal bone voxels in the coronal data set. Voxels between -100 to -400 HUare color coded in grey.Figure 2gSame as in 2e but with lower threshold expanded to -100 to -500 HU. Noteappearance of background shroud (arrow). Voxels from -100 to -500 HU are colorcoded in grey.Figure 2hSame as in 2e but with lower threshold contracted to -150 to -250.

Noteappearance of background shroud (arrow). Voxels from -100 to -500 HU are colorcoded in grey.Figure 2hSame as in 2e but with lower threshold contracted to -150 to -250. Note drop-out ofbone density voxels from the frontal bones (arrow).Figure 2iSame as in 2e but with threshold of -100. Note fill-in of soft tissue between skinsurface and bone, and drop-out of bone voxels from the frontal bones (*).203Top T....0 ' - - - ' - - - ' - - - - - ' - - ' - - ' - - - - ' - - ' - - ' - - - '0.0100.0200.01a2a1b(.) C\I.Q C\I.

MOGENEOUS METAL PARTS BY INFILTRATIONAdam M. Lorenz1, Emanuel M. Sachs1, Samuel M. Allen2, and Michael J. Cima2 1 Department of Mechanical Engineering 2 Department of Materials Science and Engineering Massachusetts Institute of Technology, Cambridge, MA 02139AbstractInfiltration of powdered metal parts made by SFF processes enables densification with negligible dimensional change, but typically uses a dissimilar infiltrant material resulting in poor corrosion resistance, machinability, and a heterogeneous composition to certification for critical applications. A new approach called transient liquid-phase infiltration is described using an infiltrant composition similar to that of the powder skeleton, but containing a melting point depressant. Upon infiltration, the liquid undergoes diffusional solidification at infiltration temperature and eventually the composition becomes homogeneous.

Upon infiltration, the liquid undergoes diffusional solidification at infiltration temperature and eventually the composition becomes homogeneous. Parts over 20 cm tall have been fabricated through careful selection of skeleton and infiltrant compositions, skeleton powder size, and infiltration technique. The work presented in this paper uses a nickel- silicon alloy to infiltrate a skeleton of pure nickel powder.inconduciveIntroductionSeveral Solid Freeform Fabrication technologies use powdered metal as the raw material to produce a net shape part skeleton that is less than 70% dense. The skeleton must then undergo further processing to achieve full density and the desired mechanical properties. The densification is typically done either by light sintering and infiltration with a lower melting temperature alloy or by high temperature sintering alone.

The densification is typically done either by light sintering and infiltration with a lower melting temperature alloy or by high temperature sintering alone. In the first method of densification, the part’s dimensional change is typically only ~1%, making it suitable for fairly large (~0.5 m on a side) parts. However, the resulting material composition is a heterogeneous mixture of the powder material and the lower melting temperature infiltrant, often with poor machinability and corrosion resistance. In critical applications such as structural, aerospace and military, a material of homogeneous composition is preferable for certification. Sintering the powder to full density will result in a homogeneous final material, but a part will undergo ~15% linear shrinkage if it starts out at 60% density.

Sintering the powder to full density will result in a homogeneous final material, but a part will undergo ~15% linear shrinkage if it starts out at 60% density. For this reason, full density sintering is typically only used for smaller (< 5 cm on a side) parts.This paper describes a process for making large, homogeneous parts from a powder skeleton using an infiltrant material similar in composition to the skeleton material, with the addition of a small amount of melting point depressant. The liquid infiltrant fills the void space in the skeleton, then undergoes diffusional solidification to achieve a homogeneous final part composition. The process is shown schematically in Figure 1 and is referred to here as Transient Liquid-phase Infiltration (TLI) due to the phase transition of the infiltrant. In some cases the diffusional solidification can prevent the infiltrant from filling a part skeleton.

In some cases the diffusional solidification can prevent the infiltrant from filling a part skeleton. Determination of conditions that allow complete infiltration is a major focus of this work.69~1 min~1 hourSintered powder skeletonLiquid infiltrant fills void spaceDiffusional solidification, homogenizationFigure 1: Transient Liquid-phase InfiltrationA nickel-based material system was chosen for development of the TLI concept because transient liquid-phase nickel brazing alloys are commonly used in the aerospace industry and nickel alloys are of interest in some DoD applications. Silicon was chosen as the melting point depressant primarily because it has a lower diffusivity than boron and phosphorous, the other common melting point depressant in nickel brazing alloys. In addition, silicon has a high enough solubility in nickel to allow complete solidification at the infiltration temperature.

In addition, silicon has a high enough solubility in nickel to allow complete solidification at the infiltration temperature. Figure 2 shows the nickel-rich side of the binary Ni−Si phase diagram1 and points out the composition of the skeleton and infiltrant as well as the resulting homogenized composition at an infiltration temperature of 1180°C. The diffusivity of the melting point depressant affects the amount of time the liquid has to fill the skeleton before it freezes. The diffusivity of Si in Ni is ~10-13 m2/s at 1200°C, resulting in a = 5 µm after 1 minute. characteristic diffusion distance of This is a sufficiently small distance to allow infiltration of nickel powders greater than 50 microns in diameter.Final part compositionInfiltrantSkeletontD ⋅4Figure 2: Ni−Si phase diagram.Study of Fluid Flow and Infiltration RateInfiltration is the result of capillary induced pressure driving the fluid into the skeleton.

In a non-reactive system, fluid will continue to flow into the skeleton until the gravitational head, pgh, balances the capillary force. The maximum capillary rise height can be calculated accordingly:hmax=1 ρ g   γLV⋅cosVp( ) θ Sp   =( γ − 16 ερ gD) ε(1)where h is fluid height, ρ is the fluid density, g is gravity, γLV is the liquid/vapor interfacial energy, θ is the contact angle of the liquid to the solid, Sp is the surface area of the pore space and Vp is the volume of the pore space.2 The right most term in the equality lumps the wetting ( )θ angle into the surface energy term , and assumes the powder bed is composed of uniform spheres of diameter D with a void fraction ε to substitute for the ratio of surface area to volume. Due to the high surface energy of most liquid metals, a skeleton of 100 µm powder will typically result in capillary rise heights greater than 1 meter.

The permeability of various powder skeletons was determined by measuring the flow rate of oil (Multitherm 503) through the skeletons at room temperature under known pressure conditions as shown in Figure 3 at right.3 Skeletons for these and all subsequent experiments were composed of hydrometallurgically processed nickel powder sintered at 1200ºC for 1 hour. Skeletons of 90−125 µm and 212−300 µm powder resulted in permeabilities of 5.0 x 10-12 m2 and 1.7 x 10-11 m2 respectively.∆ Pporous skeleton sampleFigure 3: Permeametry.The rate of infiltration into the skeleton was measured experimentally and compared to a model of the fluid flow based on Darcy’s Law. The model uses the same assumptions as Equation 1 for the capillary driving force with an extra term for the head losses due to gravity:dh dt=K εµ  ( γ − 16 ε hD) ε−ρ g  (3)Experiments were conducted at room temperature with oil and the results are shown in Figure 4.

The measured permeability predicted a slightly faster that observed experimentally, so a second curve is shown with permeability selected to match the experimental data. This permeability was 40% less than that measured by permeametry.infiltrationthanOil infiltrating 90-125 µm Skeleton400)m m(t h g e Hi300200100ExperimentModel (K=5e-12)Model (K=3e-12)Infiltration height0051015202530Time (hrs)Figure 4: Infiltration rate experiments with oil at room temperature.Measurements of the infiltration height of liquid metal inside the furnace was not directly observable, so the rate of infiltration was calculated based on mass increase. Figure 5 schematically shows the apparatus used to measure the mass of the skeleton during a transient liquid phase infiltration of a similar skeleton, along with experimental results.

Figure 5 schematically shows the apparatus used to measure the mass of the skeleton during a transient liquid phase infiltration of a similar skeleton, along with experimental results. The skeleton was suspended from a 0.010” diameter molybdenum wire and the force on the wire was measured via71a load cell and a datalogger at 0.05 second intervals. The contribution due to surface tension was subtracted from the measured force, and the resulting value was converted to infiltration height by dividing by the density of the liquid and the cross-sectional area and void fraction of the skeleton.

The model of the fluid flow is shown for the same skeleton permeability values as Figure 4, with the fluid properties of the liquid metal substituted for those of oil.Load cellNi-Si into 90-125 µm Skeleton @ 1170 °C10080)m m(60t h g e Hi40ExperimentModel (K=5e-12)20Model (K=3e-12)00306090120150Furnace hot zoneTime (sec)Figure 5: Infiltration rate experiments with molten metal and mass measurement setup.Comparison of Figures 4 and 5 demonstrates that the time scale of the liquid metal infiltration is much faster than for infiltration with oil. This difference arises because the surface tension is ~50 times greater for the liquid metal. The significant deviation of the experimental results in Figure 5 from the model prediction can be explained by the “choking” effect of the infiltrant solidifying as it flows into the part.

The significant deviation of the experimental results in Figure 5 from the model prediction can be explained by the “choking” effect of the infiltrant solidifying as it flows into the part. The next section will describe the attempt to characterize this solidification behavior.Solidification Quenching ExperimentsThe solidification rate was measured experimentally through a series of quenching experiments. A bundle of 50 µm diameter nickel wires was sintered and then the bundle was dipped into a bath of molten Ni−10Si infiltrant. After holding at the infiltration temperature for variable time periods, the samples were air quenched. The samples were then sectioned and electropolished to reveal the microstructures shown in Figure 6 below. Wire was used instead of powder so that the motion of the solidification front could be referenced to the initial wire diameter and all motion would be perpendicular to the cross-sectioned view.

Wire was used instead of powder so that the motion of the solidification front could be referenced to the initial wire diameter and all motion would be perpendicular to the cross-sectioned view. Dashed white lines have been superimposed on the images to show the initial wire diameter. The outward progression of the solidification front can be seen by comparing the pictures in sequence. The72Linear actuatorarea with a eutectic microstructure reveals the part of the sample that was liquid at the time of the quench, and is bordered by the solidification front. The black crevice that appears within the dashed white circle is an effect of the electropolishing used to reveal the microstructure. The crevice corresponds to the distance to which silicon diffused into the initially pure nickel powder. A planar diffusion couple would result in a compositional profile of an error function with characteristic diffusion distance.

A planar diffusion couple would result in a compositional profile of an error function with characteristic diffusion distance . For a diffusivity of silicon through nickel of 7.85 x 10-14 m2/s at 1180°C,4 this distance would be 1.8, 5.6 and 9.7 microns for 10, 100 and 300 seconds respectively. These expected diffusion distances correspond reasonably well with the observations in Figure 6, especially considering the difference in geometry.tD ⋅4LiquidSolidification frontInner black creviceAfter 100 sec at 1180°C Figure 6: Optical micrographs showing sequence of infiltrated bundles of 50 µm nickel wire.After 10 sec at 1180°CAfter 300 sec at 1180°CInfiltration Distance LimitsIn order to maximize the distance the infiltrant can reach before freezing in TLI, a gating mechanism is used to prevent the infiltrant from contacting the skeleton prior to infiltration. Premature contact would allow diffusion to occur and the infiltrant to freeze just as it begins to melt.

Premature contact would allow diffusion to occur and the infiltrant to freeze just as it begins to melt. Other problems could arise if liquid begins to flow before the system has reached thermodynamic equilibrium. Several gating mechanisms have been used in practice such as dipping a suspended part into the melt or initial confinement of the melt followed by opening up a path for the melt to flow out and contact the skeleton.The particle diameter of the powder affects both the fluid flow and the amount of time before the infiltrant freezes. Equation 3 describes the fluid flow behavior and because the permeability of a skeleton (K) is proportional to the square of the powder diameter the net result of using larger powder is faster liquid flow. Further, the larger powder decreases the surface area for diffusion and provides more time for liquid to flow before the porous network is closed off by solidification.

Further, the larger powder decreases the surface area for diffusion and provides more time for liquid to flow before the porous network is closed off by solidification. Figure 7 clearly shows how the infiltration distance (in this case height) is affected by the powder size of the skeleton. The capillary rise limit shown corresponds to the maximum height in Equation 1 with liquid metal properties of γ = 1.2 N/m and ρ = 8 g/cc and a void fraction of ε = 0.4. The rise limit will remain fairly consistent with different materials because many liquid metals have similar surface tension and density. The premature freezing limit shown corresponds to the experiments shown to the right using various powder sizes. Through the use of progressively larger powder size, infiltration heights as tall as 22 cm have been achieved.

Through the use of progressively larger powder size, infiltration heights as tall as 22 cm have been achieved. The freezing limit will vary significantly for different material systems because73it is affected by other properties such as the diffusivity of the melting point depressant in the skeleton, which can vary by several orders of magnitude.22 cmMaximum Infiltration Height Limits100)80m c (60t h g e Hi40Capillary Rise Premature Freezing2000100 300 200 Powder Size (microns)400Figure 7: Limitations on infiltration height.

Images at right show heights achieved with skeletons of 50−150µm, 150−300µm, and 300−400µm powder (left to right).Uniform bulk composition throughout the partDue to mass transport occurring simultaneously with the fluid transport within a part during transient liquid-phase infiltration, there exists the possibility of non-uniform bulk composition in a final part, with a higher concentration of silicon expected near the point of liquid entry. In order to prevent this variation, the liquid infiltrant can be maintained at a liquidus composition. This will ensure that any liquid present in the skeleton will be the same composition — any loss of silicon must result in solidification. Further, the liquid is already saturated with nickel and has no driving force for dissolution of the skeleton, preventing erosion.Several techniques have been used to hold the composition of the infiltrant at its equilibrium liquidus composition and prevent it from dissolving the skeleton.

Excess Ni powder was added to the melt in sufficient quantities that the bulk composition would be a mixture of saturated infiltration temperature. The supply of infiltrant was then soaked infiltration at a temperature to dissolve the excess nickel. A ceramic impeller was used to stir the melt and ensure the infiltrant supply was well mixed as the temperature The was decreased again for composition of parts infiltrated with infiltrant at its liquidus composition show no signs of erosion and have a bulk composition that does not change along the path of infiltration. This is further illustrated in Figure 8, showing a capillary infiltrated with infiltrant at itsliquid andsolid atthetemperature slightly abovethetheinfiltration.Figure 8: Uniform bulk composition along path of infiltration74liquidus composition.

The composition profile at a point just behind the infiltration front (cross- section B) would consist of pure nickel and infiltrant at its liquidus composition because very little time has elapsed to allow for diffusion. After the liquid has continued to fill the capillary, a similar location just behind the infiltration front at cross-section A’ would have an identical composition profile to B. Even though some diffusional solidification will occurr at B’ while the liquid is still infiltrating, the fixed composition of the liquid ensures that the bulk or average composition will be the same at point B, A’, and B’ — constant with time and position.Homogenization and mechanical propertiesThe time necessary for homogenization of a final part will depend strongly on the diffusivity of the melting point depressant and the geometry or powder size. Figure 9 shows cross-sections of a skeleton of ~300 µm diameter powder that was infiltrated with Ni−10Si at 1170ºC.

Figure 9 shows cross-sections of a skeleton of ~300 µm diameter powder that was infiltrated with Ni−10Si at 1170ºC. The three sections of the infiltrated skeleton shown were heat treated at 1200ºC for 1, 5 and 12 hours (left to right). Compositional analysis was conducted using an electron microprobe. The dark regions of the initial image correspond to high silicon content (>6%). After 5 hours of heat treatment, the compositional profile was nearly uniform, but regions of localized porosity are visible in the image — perhaps due to differential mass transport (Kirkendall effect). After 12 hours of heat treatment, the composition is uniform and any remaining porosity is evenly distributed.Figure 9: Homogenization of infiltrated 300µm powder after 1, 5, and 12 hours at 1200ºC.The binary Ni−Si alloy used for these experiments is not available commercially and expectations for mechanical properties were not high.

Comparison of an infiltrated specimen to a casting of identical composition showed that the transient liquid-phase infiltration process could produce a final material with similar properties. Tensile specimens were machined from an infiltrated bar homogenized for 12 hours and from a cast bar cooled at 5 ºC/min from 1400ºC. The tensile test results are shown in Figure 10 along with an image of the infiltrated specimen after failure. The elastic behavior and yield strength of both materials were nearly identical. The elongation of the infiltrated specimen was moderate at 6%, but still much less than that of the cast specimen. The mechanism causing the different behavior under high strain is not yet understood, but could be due to differences in porosity.75Tensile Test Results500) a P M(400300s s e r tS200100InfiltratedCast0051015 Strain (%)202530Figure 10: Tensile properties for infiltrated and cast specimens of the Ni − 4wt% Si.

Image at right shows the infiltrated tensile specimen after failure.SummaryTransient liquid-phase infiltration of a PM skeleton has been shown to achieve homogeneous final part composition without the dimensional change associated with full-density sintering. This densification concept is broadly applicable to any of the SFF processes that produce green PM skeletons, as well as more traditional PM processes such as Metal Injection Molding. A binary Ni−Si alloy was used to explore the important tradeoff between infiltration and diffusion rate that can prevent the liquid from filling the part due to diffusional solidification. Using infiltrant at a liquidus composition ensures uniform bulk composition throughout the part. Through gated introduction of the liquid infiltrant to the skeleton and control of powder size within the skeleton, parts greater than 20 cm tall were infiltrated to full density.

Through gated introduction of the liquid infiltrant to the skeleton and control of powder size within the skeleton, parts greater than 20 cm tall were infiltrated to full density. Subsequent heat treatment achieved a homogeneous final composition and the mechanical properties compared favorably to a cast specimen of the same composition.AcknowledgmentsThis research was sponsored by the Office of Naval Research, Contract #N00014-99-1-1090. The measurements of permeability and infiltration rate with oil were conducted by Vinay Prabhakar, a graduate student in the 3-D Printing research lab at MIT.1 Alloy Phase Diagrams. ASM Handbook Vol. 3. Metals Park, Ohio: American Society of Metals, 1992. pp 2·318 2 Scherer, George W. “Theory of Drying” J. Am. Ceram. Soc., Vol 73, No. 1, 1990. pp 3−14. 3 Heinzer, Peter J. “Permeametry.” Powder Metallurgy. ASM Handbook Vol. 7. Metals Park, Ohio: AmericanSociety of Metals, 1984. pp 263-265.4 Smithells, Colin J.

sign and Layered Manufacturingof a Lower Control Arm in Project MAXWELLRoy Johanson Noboru Kikuchi Panos PapalambrosDepartment of Mechanical Engineering & Applied Mechanics The University of Michigan Ann Arbor, MI 48109andFritz Prinz Lee WeissEngineering Design Research Center Carnegie Mellon University Pittsburgh, PA 15213Abstractbriefly describe a new methodology for the design and manufacture ofmechanical components and demonstrate the process for the design of an automobile suspension component. The methodology is a unique coupling between advanced technologies for product design and manufacture, which leads to the rapid realization of superior products. The concurrent design and analysis strategy yields information about the optimal structural layout, as well as details about the material composition. The fabrication of such designs requires unconventional manufacturing processes, such as layered manufacturing.

The fabrication of such designs requires unconventional manufacturing processes, such as layered manufacturing. Project MAXWELL, therefore, offers the possibility for the integration of layered manufacturing into the mainstream product development and fabrication process.1.INTRODUCTIONProject MAXWELL embodies a methodology that provides a rigorous technique forthe concurrent design of material composition and shape of components, and a new manufacturing process for their realization. At the University of Michigan (U-M), a new methodology has been developed for designing the form and material composition of mechanical and structural components based only on a description of the loading conditions and packaging requirements. At Carnegie Mellon University (CMU), a new manufacturing process has been developed for the free-form fabrication of parts from single or composite materials by thermal spray shape deposition.

At Carnegie Mellon University (CMU), a new manufacturing process has been developed for the free-form fabrication of parts from single or composite materials by thermal spray shape deposition. The goal of Project MAXWELL is to demonstrate the validity oLan integrated approach for the rapid realization of superior components. Such components will possess superior structural and mechanical properties, and will satisfy packaging and assembly requirements.The current application domain is in the design and manufacture of automobilecomponents. The design methodology has been applied to the design of stiffened sheet186metal/component panels, suspension components, and a variety of other structures. Ongoing work involves the extension of the methodology to other applications.In this paper, we briefly discuss the concurrent design of layout and material usingthe homogenization method.

Ongoing work involves the extension of the methodology to other applications.In this paper, we briefly discuss the concurrent design of layout and material usingthe homogenization method. After describing the Recursive Mask and Deposit (MD*) process, we then demonstrate the MAXWELL methodology for the design of a lower control ann for an automobile suspension.2. TOPOLOGY DESIGN IN MAXWELLA fundamental approach to the thermo-mechanical characterization of general composite materials was first put forth by James Clerk Maxwell (1831-1879) and was later generalized as the theory of mixtures to provide a rigorous foundation for studying the mechanics of composite materials (see, e.g. lFUN65]). Project MAXWELL aims at transforming those early ideas into engineering reality.The necessity of topological design in addition to size and shape design is widelyrecognized by structural engineers.

If topological changes are not allowed, size and shape optimization procedures can improve a design by roughly 5-150/0. Topological modifications can often yield 30-50% improvement. The homogenization method is a rigorous methodology to perform topology optimization. The problem is reformulated as a problem involving material distribution. Given an initial design domain, we discretize the domain and in troduce microscale voids throughout the structure. The optimization problem then becomes the determination of the size and orientation of the microscale voids inside the initial design domain. If an area is highly stressed, the voids in that area will tend to zero - the method will force solid material there.

If an area is highly stressed, the voids in that area will tend to zero - the method will force solid material there. By removing material completely from portions of the domain densely packed with voids, the optimum shape of the structure is identified, while its topology is determined by accounting for the number of "global" holes (see also Figure 1).Figure 1: Identification of the Shape and TopologyThis intuitive method of "sculpting" a structure is based on the theory of homogenization -- a method developed in the mid-1970's for the study of mechanics of composite materials. Applied mathematicians in France, Italy, and the former Soviet Union [LUR84, SAN80, TAR??] developed the homogenization theory to derive the constitutive equation of a composite material.Although the optimization process often yields homogeneous solids, we considerthe design domain to be a very specialized, fictitiously constructed composite material consisting of solids and voids.

In order to determine the best microstructure, we allow the design domain to include other composite materials, e.g., ones that can improve strength, toughness, vibrational characteristics, and other measures of structural performance. Such structures have not been realized in practice due to the lack of an attractive manufacturing process that delivers non-homogeneous and anisotropic materials. For exanlple, it is187impossible to create internal voids within a component by conventional NC machining. In MAXWELL, we use CMU's MD* process where the component is built up layer by layer, allowing the possibility of creating and orienting the voids as desired. Therefore, MD* allows serious consideration of these unusual and highly efficient structures for the first time.3. FABRICATION BY THE MD* PROCESSIn the MD* (recursive Mask and Deposit) process developed at CMU, parts are manufactured by successively spraying c,Toss-sectionallayers.

FABRICATION BY THE MD* PROCESSIn the MD* (recursive Mask and Deposit) process developed at CMU, parts are manufactured by successively spraying c,Toss-sectionallayers. Each layer may contain several different materials. The geometry of the part is not constrained and its shape and material composition can be changed continuously during fabrication. To create a part, its geometric model is first sliced into cross-sectional layers, typically 0.001 to 0.005 inches thick. For each material in a layer, a disposable mask is made that exposes the area where that material occurs. The mask is placed upon the top layer of the growing part shape and a robotically manipulated thermal spray gun traverses the areas exposed by the mask. Masks are made from paper stock cut with a laser.

Masks are made from paper stock cut with a laser. Several alternative strategies are feasible for creating support structures for the part as it grows, including retaining a part of the mask or spraying the support material in place after the primary materials are deposited. A more detailed description of the MD* method can be found in [WEI92]. Two features of MD* make it particulary attractive in the context of Project MAXWELL: perform selective material deposition, and the prospect of a robust process for the forming and joining of composite structures. Availability of the MD* manufacturing process provides the requisite technology for the realization of novel designs (at the macro and microscale level) generated by the homogenization method. Additionally, MD* has the potential to create dense and lan1inate structures of arbitrary geon1etric complexity, while masking also enables selective material deposition.

Additionally, MD* has the potential to create dense and lan1inate structures of arbitrary geon1etric complexity, while masking also enables selective material deposition. Therefore, different regions within a layer can be composed of different materials. Beck lBEC92] describes the manufacture of an integrated electro-mechanical structure created with the MD* process.the ability to4. EXAMPLE: LOWER C()NTR()L ARM DESIGNAs an example of the integrated design and manufacturing process of projectMAXWELL, a lower control arm was designed using the homegenization process and fabricated using MD*. This component is typical of vehicle chassis structural components; their design is typically driven by stiffness requirements. Phase I of the MAXWELL process begins with the description of an appropriate design domain and boundary conditions, shown in Figure 2. other suspension and wheel components severely limit the allowable design space.

other suspension and wheel components severely limit the allowable design space. In addition, appropriate attachment material for the strut must be provided in the final design. Translational displacements are constrained at the two pivots (to the right in Figure 1) , and loads are applied at the strut and at the ball joint. Three separate loading conditions are considered, with the primary load case being vertical (z-direction) loading at the strut attachment point.In this particular example, packaging requirements for188Figure 2: Design Domain for Lower Control ArmThe homogenization process produces a topology which possesses the maximumstiffness for a (user-supplied) constraint on the percentage of the design domain to be filled with material. Figures 3 and 4 show the result of the homogenization process, showing two cross-sections of the resultant structure.

Figures 3 and 4 show the result of the homogenization process, showing two cross-sections of the resultant structure. Note that the optimal topology which is designed using the homogenization procedure would be difficult to manufacture using traditional techniques.Figure 3: Cross-Section of Lower Control Arm at 45% Section189Figure 4: Cross-Section of Lower Control Ann at 60% SectionOften, the results from the homogenization procedure are smoothed, and detailedlocal structural constraints on allowable stresses and deflections are added and evaluated in a second sizing optimization procedure. In this example, however, we proceeded directly to the fabrication of the component using the MD* process. The data was transferred directly from the homogenization output to the MD* process through a standard stereolithography file.

The data was transferred directly from the homogenization output to the MD* process through a standard stereolithography file. The control arm was interpreted as an isotropic component composed of one material, i.e., individual elements in the structure were interpreted as either all material (solid) or all void (holes), and the control arm was fabricated from a zinc alloy.5. CURRENT STATUS AND FUTURE G()ALS ()F PR()JECT MAXWELLProject MAXWELL is a synergistic integration of two novel research efforts, one indesign and the other in manufacturing, for the purpose of establishing a sound methodology for the rapid realization of superior products. Basic research directly relevant to MAXWELL has been ongoing at the participating institutions for over four years.

The UM results to date can be summarized as the development of an integrated three-phase system for the concurrent design and analysis of superior structural components.Phase I: Based on the specified boundary conditions (type and magnitude of loads)and designable space (packaging specifications) the homogenization method !s applied to derive a gray scale representation of the material composition and distribution that is optimal relative to some desired measure of structural perforn1ance.190Phase II: Using computer vision and geometric modeling techniques, this image isinterpreted and translated into a realistic structure.Phase III: A parametric optimization model based on finite element analysis isformulated and solved to determine a complete dimensional and material description of the structure.Phase IV: The manufacture of the Phase III output (i.e., discrete parts of arbitrary geometry and possibly varying material composition) using the MD* process is performed.Phase V: The final phase in MAXWELL is the testing phase, where the Phase IV products will be subjected to various mechanical tests.

Qualitative indices of performance in Phase V will include measures such as weight to stiffness ratio, impact energy absorption rates, and fatigue life.Currently the UM system can develop structural layouts for 20, 2.50 and full 3Dcomponents. A major area of current research is appropriate geonletric representation of optimal topologies in three dinlensions. Ongoing research at CMU directly relevant to MAXWELL can be summarized as the development of the MD* process for the rapid manufacture of single or multi-material components.ACKNOWLEDGMENTSThe UM research was partially funded by the National Science Foundation. CMUhas received partial support from the National Science Foundation Engineering Research CenterREFERENCESlASH91j Ashley,S. "Rapid Prototyping SystelTlS" Mechanical Engineering, 113(4), 34 43 (1991) [BEC92] Beck, J.E., F.B. Prinz, D.P. Siewiorek, and L.E. Weiss.

W CAN MATERIAL JETTING SYSTEMS BE UPGRADED FOR MORE EFFICIENT MULTI-MATERIAL ADDITIVE MANUFACTURING?M. Baumers, C. Tuck, P. Dickens, and R. HagueAdditive Manufacturing and 3D Printing Research Group (3DPRG), Faculty of Engineering, University of Nottingham, Nottingham, NG7 2RD, UKAbstractMulti-material Additive Manufacturing (AM) platforms are able to build up components from multiple materials in a single layer-by-layer process. It is expected that this capability will enable the manufacturing of functional structures within products, such as conductive tracks or optical pathways, resulting in radically novel products with unprecedented degrees of functional density. An important variant of commercially available multi-material AM technology is material jetting, which is currently in commercial use for the manufacture of prototypes and design studies.

An important variant of commercially available multi-material AM technology is material jetting, which is currently in commercial use for the manufacture of prototypes and design studies. This paper presents a detailed process model of build-time, energy consumption and production cost for the Stratasys Objet 260 Connex system, analyzing the contemporaneous deposition of two different types of photopolymers (Veroclear RGD810 and Tangoblack FLX973).

By using this process model to anticipate the effects of various upgrades to the investigated system, such as a larger build volume and a higher deposition speed, this forward-looking paper explores pathways to enhancing the value proposition of such multi- material systems through incremental technology improvement.IntroductionSince the innovation of processes capable of additively depositing build material in the early 1980s, the spectrum of applications for these technologies has extended (Bourell et al., 2009; Shellabear and Nyrhilä, 2004; Melchels, 2012). It now ranges from the fast manufacture of design studies and prototypes to the small and even medium series manufacture of specialist high-value end use products (Lipson, 2012; Wohlers, 2012). The adoption of AM by industry carries a number of strategic implications (Cotteleer and Joyce, 2014).

The adoption of AM by industry carries a number of strategic implications (Cotteleer and Joyce, 2014). Moreover, it has been suggested that the technology has the potential for significant product innovation in the future and may alter the structure of the manufacturing sector in general (Lipson, 2012).Used synonymously with the label “3D Printing”, AM can be defined as a collection of technologies capable of “joining materials to make objects from 3D model data, usually layer upon layer, as opposed to subtractive manufacturing methodologies” (ASTM, 2012). Among AM technologies, special significance for potential product and manufacturing process innovation is attributed to processes that are capable of depositing multiple materials (Espalin et al., 2014; Diginova, 2014), allowing the embedding of functional structures to create radically novel products.It is suggested that AM technology has two main advantages over other manufacturing processes (Tuck et al., 2008).

Firstly, AM allows the manufacture of designs without many of the geometric constraints that apply to other techniques. Secondly, AM enables the manufacture of customized products in small and medium volumes at a relatively low average cost. The current state of AM technology, however, carries a set of generic processlimitations (Ruffo and Hague, 2007), acting as a barrier against the adoption of AM process in some applications:     limited material selection and characteristics, low process productivity, low dimensional accuracy, rough surface finish, repeatability and quality issues, relatively high unit cost at medium and large volumes.By constructing a detailed-analysis approach (see, for example, Di Angelo and Di Stefano, 2011) to AM process modelling, this paper assesses the build time, energy consumption and cost performance of a commercially available AM system capable of concurrently processing multiple materials, the Objet Connex 260 (Stratasys, 2014).

Using a validated process model, this paper explores how the productivity, energy efficiency and cost performance of this system could hypothetically be improved by upgrading a selection of machine characteristics and performance dimensions.The operating principle of multi-material jetting systems and the main components are described in Figure 1. Within an enclosed build volume, shown here with its cover open, photopolymer droplets are deposited by a print head (a) onto a build platform (b). Moving in the in the X / Y plane, the print head also incorporates a UV light source to initialize a polymerization reaction and a planarization mechanism to remove excess material. After finishing the deposition of material and UV exposure within a layer, the build platform indexes down by one increment in the Z direction and the deposition process for the next layer begins.

After finishing the deposition of material and UV exposure within a layer, the build platform indexes down by one increment in the Z direction and the deposition process for the next layer begins. Fresh build material is fed to the jetting head from multiple material cartridges (c), each one containing a separate build material. An additional material required for the deposition of sacrificial structures connecting parts to the build plate and to support overhangs is supplied from support material cartridges (d). The excess material removed during the build process by planarization is transferred into a waste container (e).Figure 1: Main components of a material jetting system Image source: own workThis build cycle is repeated layer-by-layer until the build operation is complete and the platform (b) can be removed by the machine operator. For additional details on the operating principle of such material jetting processes, see Gibson et al. (2010).

As with other AM techniques, it is suggested that the technical capability and cost performance of such platforms must be enhanced to increase the application spectrum for material jetting technologies to include more true manufacturing applications, (Diginova, 2014). In particular, material jetting technology carries the promise of being an important enabler for multifunctional AM.

In particular, material jetting technology carries the promise of being an important enabler for multifunctional AM. The promise of multifunctionality is that it will allow the manufacture of a new generation of high value multifunctional products, featuring embedded functionalities and performing as integrated systems rather than passive components (Espalin, et al., 2014).Showing how specific machine upgrades may result in overall cost performance improvements for material jetting systems, this paper pursues two objectives: firstly, it helps establish which directions for further technology developments may be especially worthwhile to realize material jetting systems which are better suited for true manufacturing applications.

Secondly, it aims to provide orientation to the relative magnitude of the benefits resulting from such upgrades, in terms of deposition speed, energy consumption and financial cost.MethodologyTo approach the research objective, this paper builds a detailed-analysis process model of the Objet 260 Connex system. After collecting the required empirical data by evaluating a build experiment based on a dedicated test geometry (as done, for example, by Mognol et al., 2006), the data are inserted into an existing general purpose framework for the combined estimation of build time, energy consumption and cost (see Baumers et al., 2013).

Following the validation of the process model by measuring its estimates against a real build experimentbased on parts drawn from a basket of three multi-material test specimens, the model is ready to be used to explore the effects of various upgrades to system capability.Building a detailed process modelGenerally, build time estimators form a suitable costing approach for capital-heavy production processes (see, for example, Atrill and McLaney, 1999). Thus, build time estimation provides the foundation for several AM production cost models (Alexander et al., 1998; Byun and Lee, 2006; Campbell et al., 2008). Moreover, it has been shown than some AM processes operate efficiently only where the available capacity is fully utilized (Ruffo et al., 2006; Baumers et al., 2011). This implies that to claim that an AM process is used at full capacity, it may be necessary to configure the build process to contain the maximum possible number of parts.

This implies that to claim that an AM process is used at full capacity, it may be necessary to configure the build process to contain the maximum possible number of parts. For the automated filling of AM build volumes, workspace packing algorithms have been implemented (Wodziak et al., 1994; Nyaluke et al., 1996; Ikonen et al., 1997).The process model developed for this paper is based on an adaptation of an existing general purpose framework combining an automated build volume packing technique with build time estimation (described in detail by Baumers et al., 2013). Implemented in C++ using the open source development environment Dev-C++ (v.4.9.9.2), the model is based on voxel approximations of part geometries, as proposed by Hur et al., (2001). To keep the implementation relatively concise, a number of simplifications were made.

To keep the implementation relatively concise, a number of simplifications were made. These include the following:to create a suitable build volume configuration, the model is based upon rough voxel representations of parts drawn from a basket of three multi-material demonstration components (with a resolution of 5 mm). This discretizes the problem of placing irregular and continuous geometries. As the main purpose of the voxel representations is to facilitate build volume packing, they do not contain information on the materials present in each voxel.In order to eliminate the possibility of anisotropic material properties and unpredictable part behavior, part rotation is constrained to the vertical axis. For further ease of implementation, the rotation of the test components is also limited to discrete 90° steps.On the Objet Connex 260, all parts must be connected to a removable build plate forming the build volume floor, as shown in Figure 1 (b).

Therefore, the automated packing functionality considers only arrangements in which all parts are placed on the substrate, effectively limiting part movement to the X / Y plane.As the deposition regime for multiple photopolymer materials deposited by the Objet Connex 260 system is very similar, if not identical, the time estimation part of the proposed model is not designed to distinguish between different build materials.To build the process model, the first step is to estimate build time, TBuild, which is obtained by combining data from a hierarchy of three elements of time consumption: (i) fixed time consumption per build operation, TJob, including machine warm up, (ii) total Z-height-dependent time consumption, obtained by multiplying the fixed time consumption per layer, α, by the total number of build layers l, and (iii) the total time associated with the amount of material deposited in each layer.

Equation (1) summarizes TBuild, using a triple Σ operator to expresses the summation of the time needed to process each voxel in a three-dimensional array representing the discretized build configuration:( ) ∑ ∑ ∑ ()(1)The processing of each voxel belonging to part ‘i’ is modelled by multiplying a platform specific time increment for deposition (β), essentially expressing the time needed to deposit material corresponding to a single mm2 of cross-sectional area, by the total number of such increments associated with the (5 mm)3 voxel. This measure is multiplied by the ratio of the true volume of the part, VPi (inclusive of supports), and the volume of the voxel approximation, VAi. In this model, no allowance is made for build preparation and machine cleaning. It is felt that the time spent on these activities is difficult to measure and very much at the discretion of the machine operator.

It is felt that the time spent on these activities is difficult to measure and very much at the discretion of the machine operator. It could be argued that these activities take place during the hours in which the machine is not operating.The total energy used by the build operation, EBuild, is modelled simply by adding the energy consumed to start the system up (EJob) to the time dependent element of energy consumption, which is obtained by multiplying the measured process energy consumption rate ̇ by the build time estimate TBuild :( ̇ )(2)The final element of the model is the specification of a cost estimator CBuild employing the estimator of build time (TBuild) and total energy consumption (EBuild). As the product of an activity-based costing model, CBuild is obtained by adding the total time-dependent indirect costs, obtained by multiplying an indirect cost rate ̇ by TBuild, and adding estimates of direct cost contributions in terms of raw material and energy.

The costs incurred for raw materials, of which there are three (see Table 1), are obtained by forming the dot product between a three element vector of used material volume m and a three element vector of raw material prices p. Material wastage occurring due to a planarization device built into the print head is accounted for by a uniform waste factor ω. The total energy costs are simply obtained by multiplying the energy price by the consumption estimate EBuild. Thus, CBuild can be modelled as follows:( ̇ ) ([] [] ) ( )(3)Data collection experimentAfter specifying the estimators for TBuild, EBuild and CBuild, the next step is to collect empirical data for the Object Connex 260.

As observed during previous research (Baumers et al., 2014), the layer-by-layer operating principle usually found in 3D Printing provides an opportunity to investigate whether cross-sectional area impacts the build time per layer.To facilitate this investigation, a new type of multi-material test part has been designed. It features variation in terms of cross-sectional area and material composition along the Z dimension. The resulting test part, shown in Figure 2, exhibits discrete variation in the two parameters, build material and cross-sectional area, in seven horizontal sections, each with a Z-height of 0.5 mm.

By reducing the cross-sectional area of the test part, from 5500 mm² down to 900 mm², as summarized in Table 2, it is possible to investigate if this measure of area correlates with the time needed to process individual layers, hence exploring if the deposited geometry has an effect on build speed.Figure 2: Layer-by-layer design of the test part Image source: own workTable 2: Test part characteristics, in 0.5 mm increments of Z-heightPart section (from top to bottom)Top square (30 × 30 mm)MaterialTangoBlackLayer area (mm²) 900Top square (30 × 30 mm)VeroClear900Middle square (50 × 50 mm)Mixed (52% VeroClear, 48% TangoBlack)2500Middle square (50 × 50 mm)TangoBlack2500Middle square (50 × 50 mm)VeroClear2500Bottom star shape (110 × 110 mm)TangoBlack5500Bottom star shape (110 × 110 mm)VeroClear5500Supports (~110 × ~110, height: 1.06 mm) * = Estimate based on the plan view of the test partSupport SUP7055500*To be reflective of full system capacity utilization (Baumers et al., 2013) a total of four multi-material test geometries were included in the data collection experiment, which was run once.

The empirical data on build time were extracted from the system’s internal log files and process energy consumption was monitored using a Yokogawa CW240 digital multi- purpose power meter, logging actual real power consumption in a 0.1s measurement cycle.Validation experimentTo validate the model, the build time and energy consumption of the actual machine specification were estimated and then compared to experimental measurements on the realsystem. To eliminate the possible effect of capacity underutilization (Baumers et al., 2011), a full build experiment was specified by drawing multi-material test parts from a representative basket of (non-functional) multi-material sample components. This basket, shown in Figure 3, contains a bearing block with embedded structures resembling conductive tracks (a), a belt link component with an internal structure approximating RFID functionality (b), and a small end cap with embedded identification markings (c).

This full build experiment was run once.Figure 3: Rendering of multi-material model parts, with embedded structures highlighted Image source: own workResultsData collection resultsThe data used to construct the build time model were extracted from the internal log files generated by the Objet Connex 260. The data flow into the above specification through the central parameters of the model, α1, the fixed time consumption for the deposition of each layer, and β, the time needed to deposit material corresponding to each mm2 of cross sectional area. The estimates were obtained from an ordinary least squares regression of measured layer deposition time on cross-sectional area (α1 = 21.69, β = 0.0007). It is noteworthy that the control system’s measurements of the layer time during the deposition of the first nine layers of the support structure were inconsistent (ranging from 0 s to 310 s with a mean of 45.78 s).

This indicates a control system malfunction and results in a low R2 measure of the regression (0.048).In the final model specification, which was chosen to be reflective of higher capacity utilization than the data collection experiments, the parameter α1 was replaced by a modified parameter α2, which was obtained by calculating the mean of the measured layer deposition time of the lower three sections of the multi-material test geometry (α2 = 37.25), thus effectively ignoring the narrowing of the upper sections of the test part (shown in Figure 2).Included in the information generated through the data collection experiment is a process waste factor ω, which is obtained from the control system’s own estimates for waste generated in conjunction with the experiment. Calculated simply by forming the mean of the system’s projections of material consumption and waste incurred across three materials, it is estimated at ω = 1.76.

90MJJ/s$/h*$/kg*Adapted from Baumers et al. (2013)Sys Ltd. (2014)CostMaterial cost, TangoBlack (p2)Material cost, Support (p3)419. 40142.02$/kg*$/kg*Sys Ltd. (2014)Sys Ltd. (2014)Waste factor (ω)1.76Energy price (ep)0.031$/MJ* Adapted from Baumers et al. (2013)= estimated using a $/£ exchange rate of 1.71Validation of model performanceFigure 4 shows the full build experiment, as specified by the build volume packing algorithm. The full build contains 3 bearing block components, 6 belt links, and 15 end caps. Of the available 2500 build volume floor voxels, sized (5 mm)3, the algorithm has filled 2151.

The reported estimator performance suggests a reasonable degree of accuracy.Table 4: Estimator accuracy against experimental dataNumber of parts in the buildModel estimate24Validation experiment 24ErrorTotal build time (TBuild)1406.26 min1206.85 min14.18%Total energy consumption (EBuild)44.95 MJ37.79 MJ18.93%Using the modelOnce the combined model has been assembled and tested, individual aspects of this model can be modified in a “ceteris paribus” manner (keeping everything else constant) to explore how individual upgrades to the investigated AM platform affect the overall measures of machine productivity, process energy consumption and financial cost.

It is useful to analyze the effects in terms of absolute deposition rate (cm3/h), specific energy consumption (MJ/cm3) and specific cost ($/cm3) and also in terms of relative percentage improvements over the platform’s performance in its baseline configuration.In total, 12 machine characteristics were upgraded, one at a time, requiring the implementation to be adjusted and re-executed. The improvements analyzed include: (i) aspects of machine architecture, such as build volume size and layer thickness, (ii) machine performance characteristics, such as deposition speed and machine lifespan, and (iii) reduced input prices for raw materials and energy. The idea guiding this analysis is to explore the effect of a 20% improvement in each characteristic.

It is apparent that build volume size increase (in the X, Y and Z dimensions), layer thickness increase and faster deposition speed have significant cost reducing effects (ranging from -16.63% to -11.42%). Other improvements, such as energy consumption reduction, faster machine warm up, and reduction of the price of the secondary build material (of which only approximately 6.24 cm3 were deposited), and a reduction of the energy price have only a negligible effect on the platform’s cost performance.Figure 5: Relative effect on specific cost of system improvementsImage source: own workDiscussionThe constructed model indicates that AM platforms based on material jetting technology appear to benefit significantly in terms of build rate and process cost if improvements towards higher machine capacity and throughput are made.

This scalability aspect is noted in the literature (Gibson et al., 2010) and is evidenced through experimental high volume manufacturing platforms based on direct material jetting (TNO, 2011; Project Ara, 2014).The presented model also suggests that scaling up the dimensions of the build volume produces a greater benefit in all three assessed categories (time, energy and cost) than a corresponding increase in the deposition speed. This, possibly counterintuitive, result occurs due to the specification of the build size increase as a 20% increase in the dimensions of the build volume, which does not equate to a 20% increase in the build volume cuboid. Rather, it equates to an increase of approximately 73% in volume (according to the square-cube law).

Rather, it equates to an increase of approximately 73% in volume (according to the square-cube law). While not part of this investigation, it is safe to assume that a 20% increase in the volume of the build cuboid (from approximately 12,500 cm³ to 15,000 cm³) would have had a far smaller impact on platform performance.The usefulness of current material jetting systems for manufacturing applications is limited by the range of build materials available (Diginova, 2014). Commercial direct material jetting platforms are restricted to photopolymers, such as the materials investigated in this research, and waxes (Gibson et al., 2010). Such materials can have non-standard properties that constrain their functionality, such as poor mechanical properties, materialdegradation if exposed to UV radiation, and toxicity. Moreover, as demonstrated in Table 3, the materials available for the Objet Connex 260 are currently very expensive, with prices ranging from 142.02 to 419.40 $/kg.

Moreover, as demonstrated in Table 3, the materials available for the Objet Connex 260 are currently very expensive, with prices ranging from 142.02 to 419.40 $/kg. Unsurprisingly, the overall cost performance of the investigated system is especially sensitive to the cost of the primary structural material (VeroClear in this case). This suggests that acceptable raw material prices, in particular for the bulk structural material, will be critical for the value proposition of material jetting processes in manufacturing.The general purpose framework adapted by this paper has originally been specified to be applicable to all variants of AM, with an emphasis on powder bed fusion systems (see Baumers et al., 2013). This analysis shows that the differing operating principle of material jetting technology makes the application of such a general model difficult.

This analysis shows that the differing operating principle of material jetting technology makes the application of such a general model difficult. The reason for this lies in the fact that jetting processes do not necessarily exhibit a relationship between deposition volume / geometry and build time. Such a relationship has been observed in filament extrusion systems and, to a limited extent, in powder bed fusion systems (Baumers et al., 2011). This is because material jetting systems operate through a succession of discrete print head movements in the X / Y plane, effectively depositing material in a frontier that moves in passes across the available build space.

On the Objet Connex 260, it does not appear to be of relevance for print head movement speed if material is actually deposited or not.Complicating matters, the data collection experiment has shown that the print head will not exhaust its full movement range (both in the X and Y dimensions) if the build specification does not demand it. Therefore, a relationship between the amount of material deposited and build time must be expected. Deposition in terms of print head passes can make the time needed to deposit additional material very lumpy. Effectively, the time required to deposit additional geometry depends on whether extra Y-passes, or even Z-layers, are needed. This difficulty results in an inaccurate specification using the original parameter α1, necessitating its replacement by α2.

This difficulty results in an inaccurate specification using the original parameter α1, necessitating its replacement by α2. In planned further research, which will result in a journal submission, the model proposed in this paper will be re-specified to decrease the observed estimation errors, which are currently ranging from 14.18% to 18.93%.It is noteworthy that this paper assesses a process depositing multiple materials through a multi-print head assembly moving in unison, referred to as a “print block”. This configuration implies that the process regime for the two build materials and the auxiliary support material are very similar (if not identical). In particular, this allows the first two elements of the process model (TBuild and EBuild) to be collapsed into the single material case. Clearly, this is not possible where multi-material AM is based on entirely dissimilar deposition processes such as those described by Espalin et al. (2014) or Vogeler et al.

Clearly, this is not possible where multi-material AM is based on entirely dissimilar deposition processes such as those described by Espalin et al. (2014) or Vogeler et al. (2013), for example combining filament deposition and aerosol jetting processes. Similar considerations apply to the specification of a uniform waste factor ω applying to all deposited materials. This specification may not be appropriate for platforms with more dissimilar sub- processes. This paper recommends that work should be undertaken to establish a taxonomy their of multi-material AM processes implementation.to provide structuretoinvestigations ofConclusionsThis paper has demonstrated that a combined model of build time, process energy consumption and financial cost for a multi-material AM system can serve as a very efficientavenue for initial thought experiments on how such systems could be advanced in the future.

Effectively, it shows how such a model can be used to gain an understanding of effects of specific process upgrades. In terms of modelling results, this paper suggests that increasing system productivity, build volume dimensions and the layer thickness would have a strong positive effect on process performance in terms of productivity, energy efficiency and financial cost.The importance of correctly measuring and plausibly anticipating the cost performance of novel process variants should not be underestimated. The net benefit, or value, of any technology can be investigated by analyzing the difference between the costs associated with it and its gross benefits. This implies that the potential for value creation residing within processes is inextricably bound up with their cost performance.

This implies that the potential for value creation residing within processes is inextricably bound up with their cost performance. In the near future, such patterns will be of particular interest to the proponents of multi-material AM systems which are expected to enable manufacturers and designers to efficiently create high value embedded functionality en masse. Modelling the costs of such multi-material systems credibly and showing that they can be reduced by incremental technology improvement brings the manufacturing industry one step closer to implementing this exciting and ground-breaking new technology.AcknowledgementsThe authors acknowledge the support and technical expertise provided over the course of this research by the 3DPRG’s technical staff Joe White, Mark East and Mark Hardy.References1. Alexander, P., Allen, S., and Dutta, D., 1998. Part orientation and build cost determination in layered manufacturing, Computer-Aided Design, 30 (5), pp. 343-356.2.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperHOW SIGNIFICANT IS THE COST IMPACT OF PART CONSOLIDATION WITHIN AM ADOPTION?A. Stevenson*, M. Baumers*, J. Segal*, and Sarah Macdonell**University of Nottingham, **Malvern Instruments Ltd.AbstractSuccessful implementation of advanced manufacturing technologies requires a robust pre-installation phase involving evaluation and justification of potential benefits. However, despite part consolidation (PC) being described as one of the major benefits of additive manufacturing (AM), there has been very little quantification of its potential impact on costs. This makes it difficult for organisations to consider all the benefits of AM adoption.

This makes it difficult for organisations to consider all the benefits of AM adoption. A case study research approach has been used to develop an empirical cost model based on PC for the development and production stages of a product, which can be adapted by organisations during their own pre-installation stage. The case studies involve re-design of existing sub-assemblies within a laboratory instrument producer, and the resulting cost model has been trialled using empirical data. The results show that AM has the potential to considerably reduce part count by up to 93% and associated costs by up to 85%. The significant cost saving occurs where PC results in the consolidation of numerous components thereby eliminating considerable cost elements.Introduction and backgroundAdvanced manufacturing technologies have the potential to improve manufacturing operations by reducing costs and increasing flexibility and quality (Oettmeier and Hofmann, 2016, Efstathiades et al., 2002).

However, research into implementation gained momentum when organisations were identifying that they were not delivering the required or expected benefits (Chen and Small, 1994), with 50-75% being dubbed failures in terms of flexibility, quality, and reliability (Lewis and Boyer, 2002). Implementation is defined by Voss (1988) as “the user process that leads to the successful adoption of an innovation of new technology” and generally follows their 3-phase model of pre-installation, installation and commissioning, and post-commissioning. Implementation success is defined by Udo and Ehie (1996) as “when set goals and objectives stipulated by the adoption strategy are fully realised”. Further investigations into implementation failures found that the pre-installation phase was the most critical for increasing success as it ensured a robust decision-making and planning process (Udoka and Nazemetz, 1990, Hayes and Jaikumar, 1991, Laosirihongthong et al., 2003).

Several planning and transfer models have since been developed to try and provide organisations with a structured approach for their pre-installation phase (Chen and Small, 1994, Shang and Sueyoshi, 1995, Small and Yasin, 1997, Efstathiades et al., 2002, Saberi and Yusuff, 2011). Drawing from this research, 3 main stages for pre-installation have been identified (Figure 1): evaluating potential benefits, impacts, and applications; justifying adoption of a technology or portfolio of technologies; and planning the implementation process.The nature of advanced manufacturing technologies means their impacts and benefits will be both tangible (objective and easily quantifiable) and intangible (subjective and fuzzy) (Small and Chen, 1995, Aravindan and Punniyamoorthy, 2002). Failure to properly account for all benefits during justification will mean the organisation is not assessing the true performance of the technology (Small and Chen, 1995, Small, 2006).

Pre- installation2. Installation and commissioning3. Post- installationa. Evaluationb. Justificationc. PlanningImpacts and benefitsEconomic/ financialStrategicAnalyticStrategic objectivesFigure 1 3 stages of pre-installation based on existing advanced manufacturing technology planning and transfer models and researchThough Additive Manufacturing (AM) is considered an advanced manufacturing technology, its layer-wise build method gives it capabilities far beyond traditional manufacturing processes making investment decisions highly strategic (Weller et al., 2015). Despite this and the considerable research into implementation, there has been very little transfer of pre-installation principles or techniques to AM.

Despite this and the considerable research into implementation, there has been very little transfer of pre-installation principles or techniques to AM. Some implementation frameworks have been defined (Mellor et al., 2014, Achillas et al., 2014, Deradjat and Minshall, 2015), but these assume the business case has already been made and do not provide a structured pre- installation approach for evaluation or justification. With the design freedom enabled by the layering build method, literature describes part consolidation (PC) or “integrative design” as a major benefit of AM resulting in a reduction in part count (Dietrich and Cudney, 2011, Lindemann et al., 2012, Lindemann et al., 2013, White and Lynskey, 2013, Bianchi and Åhlström, 2014, Weller et al., 2015, Yang et al., 2015, Rodrigue and Rivette, 2010). However, there is no consensus on the lifecycle impacts and very little quantification based on empirical evidence.

However, there is no consensus on the lifecycle impacts and very little quantification based on empirical evidence. When considering evaluation and justification, without understanding all the potential impacts of a technology and quantifying them where possible, a robust pre-installation process could never be performed. Current AM cost models, for example Ruffo et al. (2006), Baumers (2012), Baumers et al. (2016), Rickenbacher et al. (2013), and Piili et al. (2015), focus on well-structured process-related costs and do not systematically account for PC or wider lifecycle costs.A case study research approach has been used to develop a cost model to quantify the impacts of PC on development and production costs which can be used during the pre- installation phase of AM adoption. The model has been based on a host company, where case studies were performed to identify the potential for consolidating parts within their products in terms of the number of units reduced.

The model has been based on a host company, where case studies were performed to identify the potential for consolidating parts within their products in terms of the number of units reduced. The host company – Malvern Instruments Ltd - design and assemble laboratory instruments for material and biophysical characterisation. The seven case studies are redesigns of existing sub-assemblies that must maintain their current internal forms, performance, and assembly with higher level assemblies, with no functional or cost optimisation.

Due to the focus on the effects of PC, and to limit the scope of this study, manufacturing process cost models (build preparation to post-processing) are not constructed in this research.2552 2MethodologyDefining the cost elements impacted by part consolidationExisting literature outlined the potential impacts from PC on assembly (Atzeni and Salmi, 2012, Weller et al., 2015, Tuck et al., 2006), supply chain (Oettmeier and Hofmann, 2016, Bianchi and Åhlström, 2014, Aliakbari, 2012), and product quality from tolerance stack- up (Dietrich and Cudney, 2011, Reiher and Koch, 2016). However, very little empirical evidence was provided to prove them. Current AM cost models generally focus on well- structured process-related costs from build preparation to post-processing and do not account for any activities prior to or post manufacturing (Thomas and Gilbert, 2014). Atzeni et al.

Atzeni et al. (2010) and Atzeni and Salmi (2012) do include assembly cost based on case studies that undergo both consolidation and optimisation.The cost model in this research was based on the “Design and Development” and “Production” stages within Lindemann et al. (2013)’s activity-based costing model based on the intrinsic product lifecycle. The elements of the cost model for each stage were defined from the perspective of PC based on activities within development and production of the host company that might be impacted by the number of parts within an assembly. Process costs defined by Lindemann et al. (2013) as “machine preparation, material costs, machine costs, energy costs, and post-processing” were out of scope as these are well covered within existing cost models.

(2013) as “machine preparation, material costs, machine costs, energy costs, and post-processing” were out of scope as these are well covered within existing cost models. They also vary significantly due to machine and material costs, build speed, and post-processing requirements which would have limited the model to one type of manufacturing process.Both well-structured and ill-structured costs were identified with respect to PC, particularly when considering the impacts on product development and the supply chain (Thomas and Gilbert, 2014, Thomas, 2015, Lindemann et al., 2012). Several cost elements (activities) originally defined were identified as being difficult to quantify in terms of time due to gaps in existing literature that would not be covered by the case studies: design review, testing, assembly review, jigs and fixtures, assembly drawings, production planning, sourcing production, external assembly operations, and quality control.

Whilst reducing part count might indicate fewer design reviews for example, it is difficult to know how the new design capabilities enabled by AM might impact the time taken to review those designs. Designs may be multi-functional and more complex; but there may also be considerably fewer process limitations to consider. Other cost elements (e.g. testing, assembly, and external sub-assembly) would require information from a top-level assembly (i.e. entire product) which was not covered by the case studies in this research which are at the sub-assembly level. The final cost elements were identified as being either one-off (development) or yearly recurring (production) costs depending on their relationship with annual production volumes. Without accounting for production volumes, the model could not accurately reflect the impact of PC on production costs.

Tooling manufacture was considered within development as a one-off cost as it was an external activity for the host company and does not account for any subsequent re-tooling.Developing the part consolidation cost modelThe final cost elements were expressed in terms of a cost per unit in Table 1.2553 3Table 1 Elements of the constructed cost modelLifecycle stage Development - one-off costs independent of production volumeCost element Design and developmentTolerance analysisSpecification Development time (TDEV) multiplied by ĊLABOUR1Tolerance analysis time (TTOL) multiplied by ĊLABOUR1Allocation base Number of mechanical items (NMECHITEM) Number of mechanical parts (NMECHPART)Production - yearly recurring costs dependant on production volume NVOLUME and number of years the product will be produced PLTechnical drawings Drawings reviews Source prototypeOrder prototype Manufacture prototype Deliver prototype Tool sourcing discussion Tooling design and analysis Tooling manufacture Order productionDeliver production Book inTool settingMechanical part manufacture Kit production Standard part manufactureAssemblyTechnical drawing time (TDRAW) multiplied by ĊLABOUR1 Total review time (TREVtotal) multiplied by ĊLABOUR1 Prototype source time (TSOURCEprototype) multiplied by ĊLABOUR1 Prototype order time (CORDERprototype) multiplied by ĊLABOUR1 Cost (CPROTOTYPE)Prototype delivery cost (CDELIVERprototype) Tool discussion time (TTOOLdiscussion) multiplied by ĊLABOUR1Tool design and analysis cost (CTOOLINGd&a)Production tooling manufacture cost (CTOOLING)) Production order time (TORDERproduction) multiplied by ĊLABOUR1 Production delivery cost (CDELIVERproduction) Book in time (TBOOKin) multiplied by ĊLABOUR2 Tool setting cost (CTOOLsetting)Mechanical part manufacturing cost (CPRODUCTION)Kit production time(TKIT) multiplied by ĊLABOUR2 Standard part cost (CSTANDARD)Assembly time (TASSEMBLY) multiplied by ĊLABOUR2NMECHITEMNumber of prototypes per mechanical item (NPROTOTYPE)Number of tooled mechanical items (NTOOLITEM)NMECHPARTNumber of tooled mechanical parts (NTOOLPART)NMECHPARTNumber of standard parts (NSTANDARD) Total number of parts (NTOTAL)To facilitate understanding of the used approach, it is necessary to define several terms used in this research.

The label “Unit” is used to quantify the consolidation achieved by the case studies based on the type of part relevant to each cost element. The descriptor “Mechanical2554 4Symbol abcdefghijklmnopqrsItem” therefore refers to the number of individually designed (bespoke) parts contained within the assembly, whilst the term “Mechanical part” refers to of the number of “mechanical items” required. As tooling is not required for AM processes, there is a need to differentiate between mechanical parts that were generated via a tooled process (“Tooled mechanical items” and “parts”) as these relied on additional activities (and therefore costs) prior to production. These are counted among “Mechanical item” or “part” for all activities that are not specifically related to tooling. The term “Standard part” refers to any part which is not bespoke and are generally bought out (screws, O-rings, nuts etc.).

The term “Standard part” refers to any part which is not bespoke and are generally bought out (screws, O-rings, nuts etc.). The total number of parts included all mechanical and standard parts for calculating assembly costs. Figure 2 illustrates an example breakdown of a case study according to this classification. As process costs were out of scope, the model assumed external manufacturing of prototypes, mechanical parts, standard parts, and tooling. These and related activities were therefore expressed as a direct cost whilst the remaining elements were cost rates driven by time and labour cost.

Two distinct labour cost rates ĊLABOUR1 and ĊLABOUR2 were defined by the host company.Figure 2 Classification of Units to quantify the degree of consolidation achieved by the case studiesTo identify the cost impact of PC it is initially necessary to estimate the total development costs CDEV and the production costs CPROD using the cost elements introduced in Table 1:𝐶𝐷𝐸𝑉 = (𝑎 + 𝑐 + 𝑑 + 𝑒 + ((𝑓 + 𝑔 + ℎ)𝑁𝑃𝑅𝑂𝑇𝑂𝑇𝑌𝑃𝐸)𝑁𝑀𝐸𝐶𝐻𝐼𝑇𝐸𝑀 + 𝑏𝑁𝑀𝐸𝐶𝐻𝑃𝐴𝑅𝑇 + (𝑖 + 𝑗+ 𝑘)𝑁𝑇𝑂𝑂𝐿𝐼𝑇𝐸𝑀(1)𝐶𝑃𝑅𝑂𝐷 = 𝑁𝑉𝑂𝐿𝑈𝑀𝐸((𝑙 + 𝑚 + 𝑛 + 𝑝 + 𝑞)𝑁𝑀𝐸𝐶𝐻𝑃𝐴𝑅𝑇)𝑃𝐿 + 𝑜𝑁𝑇𝑂𝑂𝐿𝑃𝐴𝑅𝑇 + 𝑟𝑁𝑆𝑇𝐴𝑁𝐷𝐴𝑅𝐷+ 𝑠𝑁𝑇𝑂𝑇𝐴𝐿(2)Total cost for the assembly CA results from summing up development and production costs:𝐶𝐴 = 𝐶𝐷𝐸𝑉 + 𝐶𝑃𝑅𝑂𝐷(3)2555 5As can be seen from equations (1) and (2), CDEV and CPROD depend on the number of units (NMECHITEM , NMECHPART , NTOOLITEM , NTOOLPART , NSTANDARD , and NTOTAL) in each assembly.

This means that the total cost of the assembly under investigation CA must be calculated for the current unconsolidated and consolidated assembly. The potential impact of PC on costs IPC can then be calculated in terms of a potential cost saving by subtracting the total costs for the consolidated assembly CAConsolidated from the total costs for the unconsolidated assembly CAUnconsolidated:𝐼𝑃𝐶 = 𝐶𝐴𝑈𝑛𝑐𝑜𝑛𝑠𝑜𝑙𝑖𝑑𝑎𝑡𝑒𝑑 − 𝐶𝐴𝐶𝑜𝑛𝑠𝑜𝑙𝑖𝑑𝑎𝑡𝑒𝑑(4)Part consolidation case studiesCase studies were used to identify the potential for part consolidation within the host company and define the “Units” (N) of the cost model based on empirical evidence. The case studies were chosen to reflect a range of products covering different applications, production volumes, and functional requirements. Design methodologies for PC or functional integration resulting in a reduction in part count have been investigated, for example Rodrigue and Rivette (2010), and Yang et al. (2015).

Design methodologies for PC or functional integration resulting in a reduction in part count have been investigated, for example Rodrigue and Rivette (2010), and Yang et al. (2015). However, most specify stages for optimisation from a functional and/or cost perspective, which was out of scope to isolate PC from potential impacts from changes to functional performance. Existing research, knowledge, and expertise were used to redesign the case studies to reduce the number of parts, following Tang et al. (2016)’s assessment that AM can consolidate two or more components that need to move relative to eachother (i.e. integrated hinges and other operational mechanisms). The re-design process and methodology used is not a focus of this research. Certain types of component that could not be manufactured using AM processes were out of scope from consolidation: PCBs/PCAs, cables, wires, leads, mirrors, lens’, filters, glass, and motors.

Certain types of component that could not be manufactured using AM processes were out of scope from consolidation: PCBs/PCAs, cables, wires, leads, mirrors, lens’, filters, glass, and motors. Fasteners are standard parts such as O- rings, screws, dowel pins etc. were in scope but did not include those required to assemble the case studies to higher level assemblies.Trialling the part consolidation cost model on host companyThe cost and time parameters of the model were defined using data from the host company. Case studies were used to define the “units” for the consolidated and unconsolidated design, to indicate an achievable degree of consolidation and therefore calculate a realistic impact on costs. As mass customisation and agile development were not considered, it was assumed that each part was only developed once and then manufactured repeatedly as per demand.

As mass customisation and agile development were not considered, it was assumed that each part was only developed once and then manufactured repeatedly as per demand. The same cost and time parameters were used for both designs.Results and discussionIdentifying the potential for part consolidation within host company - the case studiesThe results of the case studies in terms of the number of units for the current unconsolidated and consolidated design are summarised in Table 2. Figure 3 shows two example case studies before and after consolidation. It is noteworthy that all seven case studies resulted in a reduction of mechanical parts down to one, whilst five out of seven consolidated or eliminated all standard parts. The mechanical items and parts were consolidated by taking advantage of the design freedom enabled by AM processes – which by default meant there were no mechanical items or parts requiring tooling.

The mechanical items and parts were consolidated by taking advantage of the design freedom enabled by AM processes – which by default meant there were no mechanical items or parts requiring tooling. This meant that organic forms could be easily created and tooling considerations such as split lines, rounds, and draft angles ignored. Consolidating the mechanical parts meant they no longer needed to be assembled and fixed together helping to eliminate most standard parts, whilst other were integrated to the AM part2556 6(Figure 4). Case studies 2 and 3 for example required multiple screws to assemble separately machined parts, and case study 4 required screws, O-rings, and silicone sealant to make multiple separate injection moulded parts airtight. Some standard parts were integrated to the main part where they were still required for functionality: a hinge for a lid, bearings for an enclosure door, and a cable clip for routing wires (case studies 3, 5, and 7).

Case studies 1 and 6 however still required screws for assembling a PCB and a fibre optic cable.Table 2 Results of the part consolidation case studies in terms of the number of unitsCase study Total number of items Total number of parts Mechanical itemsMechanical partsTooled mechanical items Tooled mechanical parts Standard parts1 6 Current Consolidated 2 66.67% % reduction Current 12 Consolidated 5 58.33% % reduction Current 3 Consolidated 1 66.67% % reduction Current 3 Consolidated 1 66.67% % reduction 2 Current Consolidated 0 100.00% % reduction Current 2 Consolidated 0 % reduction 100.00% Current 9 Consolidated 4 % reduction2 14 1 92.86% 20 1 95.00% 3 1 66.67% 4 1 75.00% 0 0 n/a 0 0 n/a 16 0 100.00%3 11 1 90.91% 20 1 95.00% 4 1 75.00% 4 1 75.00% 0 0 n/a 0 0 n/a 16 0 100.00%4 8 1 87.50% 24 1 95.83% 3 1 66.67% 6 1 83.33% 2 0 100.00% 4 0 100.00% 18 0 100.00%5 7 1 85.71% 9 1 88.89% 5 1 80.00% 5 1 80.00% 0 0 n/a 0 0 n/a 4 0 100.00%6 7 2 71.43% 16 2 87.50% 2 1 50.00% 2 1 50.00% 0 0 n/a 0 0 n/a 14 1 92.86%7 3 1 66.67% 3 1 66.67% 1 1 0.00% 1 1 0.00% 1 0 100.00% 1 0 100.00% 2 0 100.00%55.56%(a)(b)Figure 3 Case study 3 (a) and 4 (b) showing the before and after consolidationFigure 4 Standard parts eliminated or integrated through PC using AM2557 7Impact of PC on costs for host companyFigure 5 shows the results of the case studies in terms of a percentage cost saving achieved by the consolidated assembly for development, production, and total costs (CPC) assuming an 8yr production life.

All case studies achieved a development cost saving of between 50% and 93.5%. Case study 7 however achieved only a 10.5% production cost saving compared to between 60.7% and 85.6%, resulting in a total cost saving of only 14.4% compared to between 60.6% and 85.8%. This indicates that production costs were a bigger proportion of total costs than development. To verify, Table 3 summarises the cost saving for each cost element across all the case studies, as well as the contribution of those costs to development and production costs, and total costs. It shows that production costs accounted for significantly more than development, at 97.77% of the total cost saving. This result differ from Lindemann et al. (2013) who found that the design and development phase was the biggest contributor to total costs.

This result differ from Lindemann et al. (2013) who found that the design and development phase was the biggest contributor to total costs. However, their model was not formulated to capture the effect of PC.g n i v a st s o C %100.0% 90.0% 80.0% 70.0% 60.0% 50.0% 40.0% 30.0% 20.0% 10.0% 0.0%1234567Development93.5%75.0%75.0%93.5%80.0%50.0%86.0%Production65.7%79.0%79.0%85.6%80.7%60.7%10.5%Total over 8yr product lifetime66.4%79.0%78.6%85.8%80.7%60.6%14.4%Case studyDevelopmentProductionTotal over 8yr product lifetimeFigure 5 Bar chart showing the cost saving achieved through part consolidation for each case study for development, production, and total costs assuming an 8yr production lifeWithin Production costs, Mechanical part manufacture was the biggest contributor to the total cost saving at 66.23%.

The next most significant contributor was the cost of Standard part manufacture (11.81%), followed by Order production parts (8.10%) and then Assembly (6.97%). This reflects the results of Atzeni et al. (2010) who calculated that assembly had a smaller impact on costs than the manufacturing process, though this was based on a relatively simple 3-part assembly case study. When considering AM costs, it is well understood that these can vary significantly depending on the machine and material costs, machine productivity, and post-processing requirements (Baumers et al., 2016). This would imply that the chosen AM process for production of the consolidated assemblies would have an impact on potential cost savings from PC. As specified by AM design methodologies and research however (Yang et al., 2015, Rodrigue and Rivette, 2010), designs should be optimised from a functional and cost perspective to improve performance and reduce build time and post-processing requirements.

None of the case studies in this research have undergone any optimisation so would not give a realistic indication of manufacturing costs. This can be justified on the basis that AM costs must be considered on a case by case basis.

Figure 6 shows a simplified linear model of the degree of consolidation against production cost saving for all the case studies for each type of part. It is clear from these results that higher cost savings can be achieved when consolidating significant numbers of mechanical parts. Case study 7 does not achieve significant cost savings because it had very limited consolidation potential with only one mechanical item. However, both standard parts were eliminated or consolidated, and no tooling was required.

However, both standard parts were eliminated or consolidated, and no tooling was required. This indicates that even with no consolidation of mechanical parts, cost savings can still be achieved by eliminating tooling related costs and integrating or eliminating standard parts such as cable clips and screws.g n i v a st s o c n o i t c u d o r P %90.00% 80.00% 70.00% 60.00% 50.00% 40.00% 30.00% 20.00% 10.00% 0.00%0.00% 10.00% 20.00% 30.00% 40.00% 50.00% 60.00% 70.00% 80.00% 90.00% 100.00%Mechanical parts Linear (Mechanical parts )% Consolidation Tooled mechanical parts Linear (Tooled mechanical parts)Standard parts Linear (Standard parts )Figure 6 Consolidation against production cost saving for the mechanical parts, tooled mechanical parts, and standard parts2559 9Other tangible and intangible benefits of PC outside of development and production have not been properly considered within this research.

For example, reducing tolerance accumulation and the impacts on functionality and quality control; increasing robustness to reduce servicing and repair costs; simplifying service and end of manufacture supply chains; improving disassembly and recycling at end of life; and reducing environmental impacts. Reducing the number of parts and development time could result in reduced time to market and increased sales; whilst lower product downtime with improved functionality and reduced lead times could increase customer satisfaction. The model does also not reflect the cost of redesigning the case studies, or the costs required for the implementation process. Some AM implementation frameworks have been developed which could be used to help identify other implementation-related costs (Mellor, 2014, Deradjat and Minshall, 2015, Achillas et al., 2014, Oettmeier and Hofmann, 2016).

Existing AM cost models could also be developed to account for the potential increase in costs from increased complexity of the consolidated design, for example on removing powder from internal cavities for laser sintered components.ConclusionA cost model has been developed to quantify the impacts of PC on development and production costs. This analysis yields a simplified linear model of cost savings resulting from reducing part count which can be used to predict potential cost savings resulting from PC. It shows that the more significant the consolidation, the higher the potential cost savings. The model trial is based purely on costs of the host company, meaning the results could vary for other companies in terms of the total cost impact and the proportion of impact from each cost element.

The model trial is based purely on costs of the host company, meaning the results could vary for other companies in terms of the total cost impact and the proportion of impact from each cost element. Though parallels could be drawn from the host company in this research to an adopting organisation, this model is proposed as a tool to be used during the pre-instillation phase of implementation. It should therefore be adapted to reflect the business model, costs, and potential for PC based on case studies from the adopting organisation. The case studies in this research showcase the potential design freedom from AM processes for manufacturing complex forms, but are limited by maintaining original functionality with no optimisation.

The case studies in this research showcase the potential design freedom from AM processes for manufacturing complex forms, but are limited by maintaining original functionality with no optimisation. The model also does not reflect full lifecycle or intangible benefits of PC, and other benefits from mass customisation and functional optimisation should also be considered when justifying AM adoption.ReferencesACHILLAS, C., AIDONIS, D., IAKOVOU, E., THYMIANIDIS, M. & TZETZIS, D. 2014. A methodological framework for the inclusion of modern additive manufacturing into the production portfolio of a focused factory. Journal of Manufacturing Systems.ALIAKBARI, M. 2012. Additive Manufacturing State-of-the-Art, Capabilities, and Sample Applications with Cost Analysis. Production Engineering and Management Msc Msc, Royal Institute of Technology.ARAVINDAN, P. & PUNNIYAMOORTHY, M. 2002. Justification of advanced manufacturing technologies (AMT).

BRID ADDITIVE AND SUBTRACTIVE MANUFACTURING OF DIRECT-HEATED TOOLINGE. D. Weflen, F. E. Peters, M. C. FrankIndustrial and Manufacturing Systems Engineering, Iowa State University, Ames, IA 50011AbstractPre-heating is a common requirement for production tooling in applications such as compression and injection molding. While the carbon fiber reinforcements commonly used in large-area additive manufacturing improve the thermal conductivity of polymers, they are still far below that of metal tooling. This study presents a method for direct, local Joule heating of tooling without the need for additional heating elements. A current is induced in the composite tooling, resulting in resistance heating of the substrate. High conductivity material is locally embedded to achieve local control over the heating characteristics. Embedding of the conductive material is accomplished by selectively switching material compositions during the printing process.

Embedding of the conductive material is accomplished by selectively switching material compositions during the printing process. Demonstration tooling is produced using hybrid additive and subtractive manufacturing using an AMBIT XTRUDE in a HAAS machining center and evaluated with thermal imaging. Direct heating of tooling expands the potential applications of additive manufacturing by overcoming the challenges of low thermal conductivity materials.IntroductionProduction tooling is a proven successful application of additive manufacturing (AM) for the reduction of manufacturing lead time [1]. Material Extrusion (MEX) on a large scale has shown the potential to produce substantial polymer tooling and has been applied for molding composites and compression molding [2]. Still, polymer tooling has faced limited adoption due to a few detrimental characteristics, one of which is low thermal conductivity compared to metal tooling.

Still, polymer tooling has faced limited adoption due to a few detrimental characteristics, one of which is low thermal conductivity compared to metal tooling. The thermal conductivity of the common AM materials, even with high carbon fiber loading, is still insufficient compared to traditional metal tooling [3]. The relatively low thermal conductivity makes the traditional approach of indirect heating tooling using embedded heating elements or cartridges a poor match for this material (Figure 1a). Therefore, there is an opportunity for alternative heating methods for polymer tooling.Figure 1. Primary methods of electrical resistance heating include a) indirect resistance heating and b) direct resistance heating.1792Figure 2.

Primary methods of electrical resistance heating include a) indirect resistance heating and b) direct resistance heating.1792Figure 2. Compression mold heating approaches: a) indirect heating of a platen, b) indirect heating of tooling, c) proposed direct heating of tooling using embedded conductors, and d) proposed direct heating of tooling using external conductors.Indirect heating of compression mold tooling is the leading approach, where cartridge heaters are used to heat a platen on which the compression mold is mounted (Figure 2a). However, this approach falls short when low thermal conductivity materials are used. To overcome this, researchers have attempted to move the indirect heating elements closer to the tooling surface (Figure 2b), where precise temperature control is most critical [4]. However, this can lead to hot spots and temperature gradients throughout the tooling.

However, this can lead to hot spots and temperature gradients throughout the tooling. An alternative resistance heating approach is the direct heating of the tooling material (Figure 1b). While indirect resistance heating primarily relies on a resistive element to generate heat, it depends on the tooling material's thermal conductivity to spread that heat throughout the mold. On the other hand, direct heating relies on current flow through the material resulting in Joule heating [5]. Joule heating is more dependent on the resistivity of the material and less dependent on thermal conductivity. The carbon fiber content of polymer material can be used to tune the electrical conductivity [6], [7]. The ability to tune the electrical material properties of the tooling makes direct heating a possible approach the challenges of thermal conductivity (Figure 2c).

A machining center with hybrid additive and subtractive manufacturing capabilities is used to produce test coupons consisting of different carbon fiber contents (Figure 3). Electrically conducting elements are embedded into these samples via 3D printing, to which a voltage is applied, resulting in direct heating. The successful demonstration of direct heating of carbon fiber composite tooling shows there is potential for alternative heating approaches for tooling produced using AM.1793MethodologyTo evaluate how carbon fiber content affects electrical properties, five samples (76 x 452 x 13 mm) were printed using different carbon fiber contents (0%, 5%, 10%, 15%, 20%) by blending the neat ABS and ABS/20% CF pellets (Figure 4). Pellets were weighed and manually blended before being fed into the single screw extrusion AMBIT XTRUDE (Hybrid Manufacturing Technologies, McKinney, TX, USA).

Pellets were weighed and manually blended before being fed into the single screw extrusion AMBIT XTRUDE (Hybrid Manufacturing Technologies, McKinney, TX, USA). The ORNL slicer (Oak Ridge National Lab, Oak Ridge, TN, USA) was used to conduct process planning for the additive process, and Mastercam was used for subtractive processes.The samples with ABS/20%CF conductors were produced in six steps. Then the sample was printed using the AMBIT XTRUDE, purging the material during material changes (Figure 5a). Once the sample cooled, it was machined on exposed faces, and two 3.2 mm deep grooves spaced 50mm apart were milled with a 20° dovetail tool to help lock the next layer of material in place (Figure 5b) [8]. The grooves were filled with ABS/20% CF (Figure 5c). Once cooled, the top surface was machined flat again, and two holes were drilled in the conductors (Figure 5d). Brass threaded inserts were installed in the holes using a soldering iron (Figure 5e).

Once cooled, the top surface was machined flat again, and two holes were drilled in the conductors (Figure 5d). Brass threaded inserts were installed in the holes using a soldering iron (Figure 5e). Ring terminals attached wires to the sample, connecting it to the power supply and multimeter (Figure 5f).Figure 4. Process plan for printing test samples and the associated carbon fiber content.Figure 5. Sample production consisted of a) printing base material, b) machining slots for conductive traces, c) filling conductors via printing, d) surface machining of conductors, e) installing threaded inserts, and f) wiring into an electrical circuit.1794After testing was completed on the samples with embedded ABS/20%CF conductors, the samples were flipped over, and the second set of grooves was cut using the 20° dovetail cutting tool. A 14-gauge solid copper wire was stripped and embedded into the groove, being retained temporarily in place through thermal staking (Figure 6).

A 14-gauge solid copper wire was stripped and embedded into the groove, being retained temporarily in place through thermal staking (Figure 6). The grooves were filled with the ABS/20%CF polymer using the AMBIT XTRUDE, embedding the wires within the conductors. The ABS/20%CF polymer was used to fill the grooves on all samples produced, no matter the carbon fiber content of the base material. Instead of connecting these samples to the electrical circuit using a threaded insert, the exposed wires were connected using a lever nut splicing connector (WAGO GmbH, Minden, Germany). A complete summary of the test samples produced can be found in Table 1.Before connecting the power supply, the resistance of the circuit was measured using a Fluke 115 digital multimeter (Fluke, Everett, WA, USA). The two-wire resistance method was used.

The two-wire resistance method was used. This approach is acceptable for this preliminary study, but further work should be done to better characterize the materials' resistivity to better guide tooling systems, including integrated heating.Figure 6. Embedding copper wires in the sample.The circuit was connected to a Harrison 6284A DC power supply (Hewlett-Packard Company, Berkeley Heights, NJ, USA) set to output a measured 25V DC. The current output from the power supply was documented. The power supply had a current limiter set to 2.1A and would automatically adjust the voltage to maintain a constant current at the limit if the system exceeded that threshold. Ohm's law was used to calculate the heater's power output using the values measured. It should be noted that power output was calculated based on resistance and current at the start of the experiment and would be expected to shift with a temperature change.Table 1.

It should be noted that power output was calculated based on resistance and current at the start of the experiment and would be expected to shift with a temperature change.Table 1. Summary of the test samples and experiments conducted.Sample ID A1 A2 A3 A4 A5 B1 B2 B3 B4 B5Bulk Material ABS / 0% CF ABS / 5% CF ABS / 10% CF ABS / 15% CF ABS / 20% CF ABS / 0% CF ABS / 5% CF ABS / 10% CF ABS / 15% CF ABS / 20% CFConductor Material ABS / 20% CF ABS / 20% CF ABS / 20% CF ABS / 20% CF ABS / 20% CF Copper & ABS / 20% CF Copper & ABS / 20% CF Copper & ABS / 20% CF Copper & ABS / 20% CF Copper & ABS / 20% CF25V DC ✔ ✔ ✔ ✔ ✔ ✔ ✔ ✔ ✔ ✔118V AC - ✔ - - - - ✔ - - -1795Thermal imaging was used to measure the heating rate of the samples. The temperature at the center location and the peak temperature across the sample were monitored for a ten-minute duration using a FLIR T420bx thermal camera (FLIR Systems AB, Sweden).

The temperature at the center location and the peak temperature across the sample were monitored for a ten-minute duration using a FLIR T420bx thermal camera (FLIR Systems AB, Sweden). A video was captured throughout the duration of the experiment, with a still image captured at the ten-minute mark.To evaluate the impact of a higher voltage heating system, the samples with a 5% carbon fiber content, samples A2 and B2, were also connected 118V AC power source (Table 1). Ohm's law was used to calculate the power output. The samples were monitored for ten minutes using the FLIR thermal camera.Results and DiscussionMeasurement of the sample resistance showed that as the carbon fiber content increased, the resistance in the sample dropped (Table 1). The samples produced from neat ABS, A1 and B1, both had resistance values greater than the multimeter could measure. The addition of a copper wire in the conductor groove, represented by sample group B, reduced resistance in all cases.

The addition of a copper wire in the conductor groove, represented by sample group B, reduced resistance in all cases. The rapid drop in resistance moving from a 5% to 10% carbon fiber content suggests that the percolation threshold was reached, which aligns with expectations [6]. There has been considerable work completed in the modeling and experimental evaluation of the electrical and thermal conductivity of polymer materials with conductive fillers, which can help guide future efforts to tune in the appropriate resistance for tooling [7], [9].Recorded current values from the power supply are documented in Table 1. Three samples reached the current limit, A5, B4, and B5. The system automatically drops the voltage to maintain this current limit. The measured resistance and current were used to calculate the power output for these samples (Table 1). These results suggest that if appropriately tuned, a heater could be designed to utilize the total power output of the power supply.

These results suggest that if appropriately tuned, a heater could be designed to utilize the total power output of the power supply. The ability to heat the tooling more quickly would be driven by the available capacity of the power supply and not a limitation of the tooling material properties. This heating rate can outperform tooling using indirect heating, which has the rate of heating limited by the material thermal conductivity.Table 2.

The localized heating near the terminals only occurred in the samples with the ABS/20%CF conductors. Applying Kirchhoff's Law would show that this is likely due to a voltage drop along our conductor due to the relatively high resistivity of the composite- filled polymer conductor. These samples experienced local heating to an extent where melting occurred in these locations. The melting prevented the ten-minute thermal imaging experiments from being completed on these samples.Figure 7. Thermal image of sample A5.Samples B1-B5 presented less concentrated heating than the samples that did not contain the embedded copper conductor. Thermal images after ten minutes of heating show that samples B2, B3, and B4 did experience heating along the length where copper wire and ABS/20%CF are embedded in the part (Figure 8c- e). The upper left portion of the image shows the temperature at the center of the sample, while the value in the upper right reads the maximum temperature in the sample.

The upper left portion of the image shows the temperature at the center of the sample, while the value in the upper right reads the maximum temperature in the sample. As our calculated power output predicted, sample A1 did not exhibit any heating (Figure 8a). Sample B2 shows some signs of heating, which is generally well spread throughout the part but represents only a 2.5° C increase from the ambient lab environment of 25° C. The resistance in sample B2 is likely too high to result in substantial heating at 25V, which led to further testing of that sample at a higher voltage to determine if the distributed heating would remain with more heat generation resulting from the higher current flow.Figure 8.

Thermal images of test samples after ten minutes of heating: a) sample B1 shows no heating, b) sample B2 shows signs of even heating, c) sample B3 shows concentrated heating, d) sample B4 shows concentrated heating, e) sample B5 shows concentrated heating.1797A high voltage of 118V AC was connected to samples A2 and B2. Sample A2 shows concentrated heating, but this time it is at the opposite end of the sample from the conductors (Figure 9a). This suggests that while the conductor of sufficiently low resistivity, the bulk ABS/5%CF material may have experienced inconsistent blending. This is likely, due to the short single screw extrusion system used, and the simple blending of neat ABS and ABS/20%CF pellets. Sample B2 shows similar distributed heating as observed at 25V DC, but the temperature increase after the ten-minute heating period was much greater (Figure 9b).

Sample B2 shows similar distributed heating as observed at 25V DC, but the temperature increase after the ten-minute heating period was much greater (Figure 9b). Sample B2 experienced a peak temperature increase of 42.2° C from the 25° ambient lab temperature at 118V Ac compared to only 2.5° C temperature increase at 25V DC. Importantly, the heating of samples A2 and B2 was achieved after only a ten minute period, compared to five hours to achieve similar temperatures when indirect heating was used to heat additively manufactured composite compression mold tooling [3]. While both samples A2 and B2 show some inconsistency in heating across the samples, they suggest that direct heating has potential if the carbon fibers can be distributed more evenly through the polymer matrix, and the fiber content can be tuned to the application.Figure 9.

Thermal image after ten minutes of heating for samples containing 5% carbon fiber by weight with a) ABS/20%CF conductor (A2), and b) copper wire and ABS/20%CF conductor (B2).ConclusionsDirect heating of carbon fiber reinforced polymer tooling produced using a large-scale MEX was demonstrated as having potential as an alternative approach to the indirect heating method employed today. Small sections of composite tooling were printed with various carbon fiber contents and either an ABS/20%CF embedded conductor or an ABS/20%CF and copper wire conductor. Testing at lower voltages resulted in hot spots around the conducting regions, while higher voltages resulted in more even heating. While further work is required to implement direct heating on production tooling, the results demonstrate the potential for direct heating of AM tooling.

While further work is required to implement direct heating on production tooling, the results demonstrate the potential for direct heating of AM tooling. A tremendous advantage of the proposed method is that we can avoid interrupting a printing process to have technician manually embed heating elements or conductive wires. There is even potential to eliminate the machining process for the conductive traces and simply leave a bead-width void wherever a conductive trace is needed and then switch materials in process. The only manual step would be to connect the circuit leads upon completion.1798Additional work establishing the electrical materials properties of the polymers is needed. It is likely that insufficient blending occurred between the neat ABS and the ABS with 20% CF content within the short screw single screw extruder used in this study. Better could improve the consistency of heating within the samples.

Better could improve the consistency of heating within the samples. Direct measurement of the electrical material properties independent of the demonstration assembly would also help guide the design of different heating configurations. Different carbon fiber loading in the polymer will likely result in changes to the coefficient of thermal expansion, which could lead to delamination between the dissimilar material regions after thermal cycling. However, the application of functionally graded materials may provide a potential route to managing these stresses. In addition, a strategy for placing regions of the conductor, resistor, and insulator in the tooling will be needed to achieve the required heating safely. Current could potentially be applied across the entire tool if local heating control is not required, eliminating thermal gradients in the tool.

Current could potentially be applied across the entire tool if local heating control is not required, eliminating thermal gradients in the tool. Despite these challenges, direct heating can expand the applications of large-scale MEX into tooling applications where rapid heating and precise temperature control are required.References[1] B. K. Post, R. F. Lind, P. D. Lloyd, V. Kunc, J. M. Linhal, and L. J. Love, "TheEconomics of Big Area Additive Manufacturing," 2016 International Solid Freeform Fabrication Sympossium, Aug. 2016.[2] C. E. Duty, T. Drye, and A. Franc, "Material development for tooling applications using[3][4]big area additive manufacturing ( BAAM )," ORNL Tech. Rep. ORNL/TM-2015/78, 2015. P. Yeole, S. Kim, A. A. Hassen, V. Kumar, V. Kunc, and U. Vaidya, "Large-scale additive manufacturing tooling for extrusion-compression molds," Addit. Manuf. Lett., vol. 1, p. 100007, Dec. 2021. E. Barocio, R. B. Pipes, B. Brenken, A. J. Favaloro, and N. M.

Vaidya, "Large-scale additive manufacturing tooling for extrusion-compression molds," Addit. Manuf. Lett., vol. 1, p. 100007, Dec. 2021. E. Barocio, R. B. Pipes, B. Brenken, A. J. Favaloro, and N. M. DeNardo, "Methods and apparatus for embedding heating circuits into articles made by additive manufacturing and articles made therefrom," United States Patent 10,889,056, Jan. 2021.[5] G. Sotiriadis, N. Athanasopoulos, and V. Kostopoulos, "A Study on the Effect of Joule-Heating During the Liquid Composite Molding (LCM) Process and on the Curing of CFRP Composite Laminates." The 10th International Conference on Flow Processes in Composite Materials, 2010.[6] A. Motaghi, A. Hrymak, and G. H. Motlagh, "Electrical conductivity and percolationthreshold of hybrid carbon/polymer composites," J. Appl. Polym. Sci., vol. 132, no. 13, p. 41744, Apr. 2015.[7] Q. Zhao et al., "Review on the Electrical Resistance/Conductivity of Carbon Fiber[8]Reinforced Polymer," Appl. Sci. 2019, Vol.

By solving a standard cantilever problem, we show that the hybrid method oﬀers improved performance over the standard BESO method. It is proposed that the new method is more suitable for optimizing structures for AM in a computational eﬃcient manner.IntroductionTopology optimization (TO) is the structural optimization method least sensitive to the initial design and so has the greatest capacity for realizing improved design for additive manufacture (AM). It began when Mitchell [1] established principles for the TO of truss-like structures and later, work by Rozvany and Kirsch [2, 3] extended this to allow the optimization of grillages and bar systems. Bendsoe and Kikuchi’s [4] homogenization technique signiﬁcantly advanced TO as it reduced the optimization of con- tinuum structures to that of determining the optimum parameters of holes in unit cells that constituted such structures.

Rozvany and Zhou [5] further simpliﬁed TO by introducing the SIMP (Solid Isotropic Microstructure with Penalization) algorithm to obtain practical designs. Steven and Xie [6] also proposed a simple approach to TO, called evolutionary structural optimization (ESO), though Zhou and Rozvany [7] attribute much of its ground work to Mattheck et al. [8] who used a similar procedure in adaptive biological growth. Querin et al. [9] enhanced the eﬃciency of the ESO algorithm with the bi-directional evolutionary structural optimization (BESO) algorithm. Huang and Xie [10] improved on this to solve mesh dependency and non-convergent problems. Stochastic optimization algorithms have also been used for TO. Sandgren et al. [11] and Chapman et al. [12, 13] performed genetic algorithm based TO. Kaveh et al. [14], and Luh and Lin [15] used ant colony optimization algorithms. Luh et al. [16] applied particle swarm optimization while Bureerat and Limtragool [17] preferred simulated annealing.

Kaveh et al. [14], and Luh and Lin [15] used ant colony optimization algorithms. Luh et al. [16] applied particle swarm optimization while Bureerat and Limtragool [17] preferred simulated annealing. Hybrids have also emerged, Zuo et al. [18] combined BESO and a genetic algorithm and Garcia-Lopez [19] coupled SIMP with simulated annealing. These stochastic algorithms operate on a population of solutions for a complete exploration of the design space. However, computational cost increases signiﬁcantly when they are used to solve complicated three dimensional problems. Similar to non-stochastic algorithms, stochastic algorithms rely on results from a ﬁnite element analysis (FEA) to move an initial topology to an optimum.FEA is a numerical technique used to estimate solutions for problems governed by partial diﬀerential equations [20]. It assumes a piecewise form of the equation for a given domain.

It assumes a piecewise form of the equation for a given domain. An assembly of each piece or element constitutes a mesh improved either by h−, p− or r− reﬁnement [21]. A common ap- proach to TO involves prior preparation and improvement of the mesh to an acceptable quality, with the mesh remaining ﬁxed for the duration of the TO. Authors illustrate the practicality of their TO algorithm by solving simple problems with this sort of meshing strategy. For three dimensional problems with complicated geometry, satisfactory mesh quality can only be achieved with a high number of ele- ments. Following TO, post processing operations are often a necessity owing to the characteristic rough boundaries of optima. Again, using a ﬁner mesh would reduce the extent of the roughness, however the computational resource required for this often makes them undesirable, especially for three dimensionalproblems. Both Kim et al.

Both Kim et al. [23] developed algorithms that achieve smoother boundaries at reasonable mesh sizes. However, similar to other TO, their reliance on a ﬁxed mesh limit the appearance of intricate features that could potentially enhance structural performance.[22] and Mariano et al.Until recently, the appearance of such intricate features was not desirable due to manufacturing con- straints. These small feature were penalized either during or after a TO to allow manufacturability. Recent developments in the ﬁeld of additive manufacturing (AM) have enabled a signiﬁcant increase in the level of geometric complexity achievable without a signiﬁcant increase in the cost of manufacture. This is due to the inherent layer-by-layer approach to manufacturing which eliminates the need for tooling or ﬁxtures [24, 25]. According to Hague et al. [26] “the advent of additive manufacturing will have profound implications for the way in which designers work.”.

According to Hague et al. [26] “the advent of additive manufacturing will have profound implications for the way in which designers work.”. Rosen [27] commenced a domain exploration based on sizing optimization of cellular materials called mesostructures. Interestingly, the versatility of this exploration could be enhanced by coupling an adaptive meshing strategy (AMS) with TO algorithms. In this paper, we couple the BESO algorithm to an AMS to solve a two dimensional problem. The discrete- ness of intermediate topologies in the BESO algorithm makes it suitable for our strategy. To demonstrate the performance of our method, a cantilever plate problem was solved to minimize total strain energy, C, for a given area fraction constraint, A∗.

To demonstrate the performance of our method, a cantilever plate problem was solved to minimize total strain energy, C, for a given area fraction constraint, A∗. This work is an aspect of a project titled “Topological optimized additively manufactured structural metallic components” funded by the engineering and physical science research council (EPSRC) and the innovative manufacturing and construction research centre (IMCRC).MethodCertain aspects of Huang and Xie’s [10] BESO algorithm were modiﬁed to allow an eﬀective coupling of the AMS with the BESO algorithm. This hybrid method was then based on an iterative improvement of an unstructured triangular mesh by reﬁning elements at boundaries of intermediate topologies and coarsening other regions. Using Engwirda’s [28] code, elements were selectively reﬁned while coarsening was achieved by edges collapse. This necessitated the development of a logic for recognition of elements to reﬁne and edge to collapse.

This necessitated the development of a logic for recognition of elements to reﬁne and edge to collapse. A consequence of this strategy was the degradation of mesh quality minimized by subjecting the emergent mesh to Laplacian smoothening [31] (an r−reﬁnement method) method. To illustrate performance, we solve a cantilever plate problem with area, Ai, at iteration, i, experiencing load, F , with an objective to minimize C given A∗.

To illustrate performance, we solve a cantilever plate problem with area, Ai, at iteration, i, experiencing load, F , with an objective to minimize C given A∗. Expressed mathematically as,min : C =1 2Ftu,subject to :Ai A0= A∗where u is a displacement vector.Modiﬁed BESOThe BESO algorithm was modiﬁed to involve,FEA, • Computation of elemental strain energies, λa, • Calculation of ﬁltered elemental strain energy densities, χa, • Calculation of a target area, A1, • Deletion and addition of elements based on thresholds, χth • Adaptive mesh improvement, • Comparison of change in strain energy, ∆C, against a tolerance, θ, • Iterations of steps 1 to 5 until ∆C is lower than θ, del, χth add(1), We implemented this BESO version by passing a discretized design domain, ﬁlter radius factor, R was a scalar multiple used to include contributions and evolution rate, v, to the algorithm.

The R of neighboring nodes to the ﬁltered strain energy densities of an element being ﬁltered. The v was the iterative rate of change of area of the undeleted elements. An explicit description of their use is given later. The ﬁrst step, FEA, was performed using MSC Nastran(cid:114) (MSC Software, California, Santa Ana). Elemental strain energies, λa, were outputted from MSC Nastran(cid:114) and inputted into MATLAB(cid:114) (Mathworks, Natick, Massachusetts) where subsequent steps of the BESO were implemented. We then computed elemental strain energy densities, χa, from λa through(cid:48)χa =λa Aawhere Aa was the area of element, a. Filtering χa suppressed the appearance of checkerboard patterns, and was achieved in two stages. The ﬁrst involved a distribution of χa over connecting nodes to obtain nodal sensitivities, ηb, byN (cid:88)Aaχaηb =a=1 N (cid:88)Aaa=1where N was the last element connected to node b.

The ﬁrst involved a distribution of χa over connecting nodes to obtain nodal sensitivities, ηb, byN (cid:88)Aaχaηb =a=1 N (cid:88)Aaa=1where N was the last element connected to node b. Secondly, calculating ﬁltered elemental strain energy densities, χa, from ηb and R(cid:48)through,N (cid:88)ψ(dab)ηbχa =b=1 N (cid:88)ψ(dab)b=1 where ψ(dab) = R(cid:48)y − dab, y was the radius of a circle that circumscribed the element, a and dab was the distance between node b and the center of element a. Next, a target area, A1, was computed from v according toA1 =(cid:40)Ai(1 − v), if A∗A0, ifAi A0 Ai A0> A∗ ≤ A∗A set of areas, (cid:98)Ω, belonging to undeleted elements was sorted in descending order of χa.

This order allowed the construction of a cumulative area set, (cid:98)κ, for each element through,(cid:98)κ(a) =a (cid:88)(cid:98)Ω(a)j=1The χa of an element whose corresponding cumulative area, (cid:98)κa was nearest to A1 was assigned as a threshold, χth del were deleted. The was achieved by re-associating a predeﬁned void property to them with an insigniﬁcant Young’s modulus (1% of that of undeleted elements). A similar process was then applied to pre-existing deleted elements. The χa of the element whose cumulative area, (cid:98)Ω(a), approximately equaled A1 − Ai was assigned to a diﬀerent threshold, χth add. Elements with χa above this threshold were reassociated to the existing property for undeleted elements. The mesh was then improved through the AMS detailed in the next section.del. Undeleted elements with χa below χthAdaptive MeshingThe AMS was basically composed of selecting, reﬁning, coarsening and smoothening.

Undeleted elements with χa below χthAdaptive MeshingThe AMS was basically composed of selecting, reﬁning, coarsening and smoothening. We deﬁne thefollowing sets to enhance the description of the AMS:(cid:48)(2)(3)(4)(5)(6)(cid:98)U = {All elements in design domain}, • (cid:98)Q = {Undeleted elements connected to boundary edges}, • (cid:98)A = {Elements in (cid:98)Q whose A exceed a lower limit, Amin }, • (cid:98)R = {Elements to reﬁne}, • (cid:98)N = {New elements derived from (cid:98)R}, • (cid:98)O = {Old elements,} • (cid:98)C = {Elements to delete by collapsing (thereby coarsening the mesh).} • (cid:98)X = {Points whose X Cartesian coordinates lie on the design domain boundary}, • (cid:98)Y = {Points whose Y Cartesian coordinates lie on the design domain boundary},The ﬁrst stage in the AMS was to select the elements at the edeges of the current topology that would be subject to re-meshing.

A connectivity matrix, L, was constructed to contain the total number of elements in the mesh, h. Each row, a in L held the node identiﬁcation numbers, b, of nodes connected to element, a. An edge matrix E with 3h rows and 2 columns was then constructed from L throughEi,k = Lx,ywhere,x =(cid:40)i − (cid:98) i hh (cid:99), if , ifi − (cid:98) i i − (cid:98) ih (cid:99) (cid:54)= 0 h (cid:99) = 0, y =(cid:40)(cid:100) i h (cid:101) + (k − 1), if , if1(cid:100) i h (cid:101) < 3 (cid:100) i h (cid:101) = 3, 1 ≤ i ≤ 3h, and 1 ≤ k ≤ 2Each row in E held node numbers connected to each edge. Subsets occurring only once represent edges lying on the boundary of the domain. These subsets were placed in a matrix E1. The set of elements, (cid:98)Q, connected to these edges was then determined by(cid:98)Q = {a : Ej 1 ⊆ La,(1,2,3)}where j is an integer used to select edges in E1 for each element, a.

The set of elements, (cid:98)Q, connected to these edges was then determined by(cid:98)Q = {a : Ej 1 ⊆ La,(1,2,3)}where j is an integer used to select edges in E1 for each element, a. The size of (cid:98)Q was further reduced by subjecting it to two further criteria. The ﬁrst involved a lower area limit, Amin, imposed on members of (cid:98)Q so that only elements with area, A, above Amin were passed to a new set (cid:98)R. Secondly, (cid:98)R was reduced by removing elements whose Aa falls below a certain percentile (90%). These two steps prevented an inﬁnite growth in the number of elements.As mentioned earlier, Engwirda’s [28] method was employed for reﬁnement purposes. This replaced elements in (cid:98)R with new smaller elements based on two reﬁnement templates. This was achieved by ﬁrst storing edges of elements to reﬁne in E1. The mid point of these edges were computed and including in the mesh.

This was achieved by ﬁrst storing edges of elements to reﬁne in E1. The mid point of these edges were computed and including in the mesh. Elements with all three edges belonging to E1 were divided into four elements, while elements with a single edge in E1 were divided into two on the edge found in E1. A much detailed description of this algorithm can be found in [28]. Elemental properties of reﬁned elements were inherited by corresponding new elements. After reﬁnement, (cid:98)O was constructed from (cid:98)N(cid:98)O = (cid:98)N c = (cid:98)U \ (cid:98)NConstruction of (cid:98)C was accomplished by extracting the smallest 1% of elements in (cid:98)O. This lower range was appropriate to prevent excessive distortion of elements in the mesh. Also, elements lying on the boundaries of the design domain were removed from (cid:98)C. The (cid:98)C and the mesh were then passed to the coarsening subroutine where the element collapsing [30] operation was performed.

The (cid:98)C and the mesh were then passed to the coarsening subroutine where the element collapsing [30] operation was performed. The ﬂow chart for this subroutine is shown in Fig. 1. Set (cid:98)C was iteratively checked until it was empty which initiated the termination of the coarsening operation. However, a populated (cid:98)C caused a move into a loop where a restoration point was initiated by duplicating the current mesh into Nmesh. The nth edge of the ﬁrst element in (cid:98)C was then collapsed in Nmesh by setting both the ﬁrst and second nodes of the edge to the mid point, Pmid, of that edge. This operation caused a degeneration in elements connected to the edge(7)(8)(9)Figure 1: Flow chart for coarsening subroutine based on element edge collapse.and an enlargement of other elements connected to just one node of same edge. Degraded elements were then simply removed from Nmesh without compromising it’s topology.

Degraded elements were then simply removed from Nmesh without compromising it’s topology. The quality, Qa, of elements in the vicinity of the deleted edge was evaluated according to Bhatia and Lawrence [29] metrics expressed as√ 4 l2 1 + l2 where l1, l2, l3 are the element edge lengths and 0 ≤ Qa ≤ 1. For degenerate elements Qa approaches a value of zero while regular elements were characterized by a Qa value of one. The variable mesh was replaced by Nmesh only if low quality elements (Qa < 0.01) were absent from Nmesh. Otherwise, a move is forced to the next edge and the collapsing process repeated. An element was removed from (cid:98)C once col- lapsed or if no collapsing operation for its three edges resulted in adjacent elements of acceptable quality. Further improvement in the quality of mesh was achieved by Laplacian smoothening [31]. Convergence of the Laplacian smoothening algorithm was based on the quality metric stated earlier with a minimum quality of 0.05.

Convergence of the Laplacian smoothening algorithm was based on the quality metric stated earlier with a minimum quality of 0.05. The mesh was passed back to BESO after each improvement and convergence tested by calculating ∆C3Aa 2 + l2 3Qa =|T (cid:88)(Ci−k+1 − Ci−T −k+1)|∆C =k=1|T (cid:88)Ci−k+1|k=1where T = 5, k is a counter varying from 1 to T , used to selectively pick C for the last T iterations. Variable ∆C was compared against θ to either terminate or repeat the BESO algorithm. θ assumed a value of 1 × 10−3. We hypothesize that coupling our AMS with the BESO algorithm would provide greater eﬃciency, reduced sensitivity to the starting mesh and smoother boundaries thereby reducing subsequent geometric post processing.

We proceed to give a detailed account of experiments conducted to investigate these hypotheses by solving a cantilever plate problem to minimize C.(10)(11)Test ProblemTo test the performance of the AMS, we solved a cantilever plate problem (Fig. 2) with an objective to minimize C subject to an area fraction constraint, A∗ of 0.5. A Young’s modulus of 100GPa and Poisson’s ratio of 0.3 was assumed.Figure 2: Cantilever plate subject to a point load, F , and constrained on the left side.We investigated the validity of our hypotheses with two diﬀerence set of experiments. Using diﬀerent mesh sizes we compared our strategy with the original BESO [10]. The R and v were ﬁxed at 1.5 and 1% respectively for all experiments. It was anticipated that by selecting these values for the optimization parameters, penalization of ﬁner features would be minimized. Experiments were performed on a desktop computer with Intel(R) Core(TM) CPU, 3.20GHz and 3.24GB of RAM.

Experiments were performed on a desktop computer with Intel(R) Core(TM) CPU, 3.20GHz and 3.24GB of RAM. Where the results are labelled ’BESO’, this refers to Huang and Xie’s version [10].(cid:48)Test of EﬃciencyFour experiments were performed to ascertain the computational eﬃciency for both the AMS and BESO algorithm (table 1). The starting mesh size and area limit, Amin, were varied across these ex- periments. An initial value of 0.1mm2 was assumed for Amin, while solving the problem with the AMS algorithm with 2000 elements initially in the design domain (Exp.1). A corresponding uniform mesh was created with elemental areas equal to 0.1mm2 (40,000 elements) was utilized in Exp. 2 where the BESO algorithm was applied. Two other experiments were performed using a lower value of Amin (0.02mm2). Starting with a mesh populated with 40,000 elements, the ﬁrst experiment was repeated (Exp. 3). The corresponding uniform mesh for this value of Amin amounted to 200,000 elements in Exp.

Starting with a mesh populated with 40,000 elements, the ﬁrst experiment was repeated (Exp. 3). The corresponding uniform mesh for this value of Amin amounted to 200,000 elements in Exp. 4 where the BESO was applied. Optima truss-like topologies for these four experiments are shown in Fig. 3.Table 1: Optimization run details and results for Exps. 1 to 4 comparing CPU time used, iterations to convergence and converged strain energy, C∞ and stiﬀness, K∞.Exp. Algorithm Start No. elems. Amin1 2 3 4AMS BESO AMS BESO2,000 40,000 40,000 200,000(mm2) 0.1 0.1 0.02 0.02Time (hr:min) 0:18 0:39 1:42 9:39No. iterations C∞75 78 77 78(Nmm) 2.21 2.21 2.25 2.28K∞ (N/mm) 27.11 26.85 27.03 27.00It can be seen that the boundaries of the topologies shown in Figs. 3a and 3c appear darker than the interior regions. A magniﬁed view B shown in these Figs. allows a better visualization of the mesh.

3a and 3c appear darker than the interior regions. A magniﬁed view B shown in these Figs. allows a better visualization of the mesh. This is due to the higher density of smaller elements at these boundaries, in contrast to an even shade(a)(b)(c)(d)Figure 3: Optima topologies and a magniﬁed view B for (a) Exp.1 (b) Exp. 2 (c) Exp. 3 (d) Exp. 4observed in Figs. 3b and 3d due to a near constant element size. We characterize each experiment by the CPU time consumed, number of iterations to convergence and the asymptotic strain energy, C∞. Results for these measures are contained in table 1.Exps. 1 and 2 converged to the same C∞, however Exp. 2 took twice as long as Exp. 1, also requiring a greater number of iterations. A similar trend is repeated for Exp. 3 and Exp. 4, though Exp. 3 converged approximately ﬁve times faster than Exp. 4. Also while Exp. 1 and 2 attained the same C∞, Exp. 3 and 4 diﬀered slightly. Similarity in C∞ for Exp.

3 and Exp. 4, though Exp. 3 converged approximately ﬁve times faster than Exp. 4. Also while Exp. 1 and 2 attained the same C∞, Exp. 3 and 4 diﬀered slightly. Similarity in C∞ for Exp. 1 and 2 can be attributed to approximately the same number of nodes at the end of both Exps. as shown in Fig. 4a.The number of nodes increased in Exp. 1 as the TO progressed to end with a similar number of nodes to Exp. 2. However, Exp. 3 and Exp. 4 have a signiﬁcantly diﬀerent number of nodes at the last iteration as shown in Fig. 4b. The diﬀerent number of degrees of freedom makes it diﬃculty to fairly compare these C∞ values since diﬀerent levels of FEA errors will exist in the calculation of C. Distortions at the loaded node compound this problem, necessitating a post analysis. Since the main aim of a strain energy minimization problem is to minimize deﬂection, the post analysis was performed to determining the stiﬀness of the topologies.

Since the main aim of a strain energy minimization problem is to minimize deﬂection, the post analysis was performed to determining the stiﬀness of the topologies. Stiﬀness in this context refers to the ratio between load F and the magnitude of a nodal displacement. The chosen node was in the vicinity of the loaded node, however far enough to avoid FEA errors caused by distortion. The location of this node was ﬁxed for all topologies.Boundaries of topologies were ﬁrst extracted, followed by a smoothening process to eliminate the rough boundaries without violating V ∗. All, topologies were then remeshed with second order elements and then analyzed. The number of elements was increased until a converged stiﬀness, K∞ was reached. The K∞ for the four topologies are shown in table 1. The K∞ for Exp. 1 was slightly greater (0.01%) than that of Exp. 2. Also, K∞ of Exp. 3 was slightly greater (0.001%) than that of Exp. 4.

The K∞ for the four topologies are shown in table 1. The K∞ for Exp. 1 was slightly greater (0.01%) than that of Exp. 2. Also, K∞ of Exp. 3 was slightly greater (0.001%) than that of Exp. 4. This suggested topologies achieved by the AMS were marginally stiﬀer than those using the BESO algorithm. While it could be argued that the K∞ values of Exp. 1 to 4 were approximately equal, the computational eﬃciency oﬀered by the AMS makes it attractive for TO for AM.(a)(b)Figure 4: Plots of C and nodes against i for (a) Exps. 1 and 2 (b) Exps. 3 and 4Sensitivity TestA second set of experiments was performed to compare the sensitivity of the AMS to the starting mesh compared to that of the original BESO. Eight experiments were formulated for this purpose (table 2), the ﬁrst four were performed with the AMS and the second four using the BESO algorithm.

Eight experiments were formulated for this purpose (table 2), the ﬁrst four were performed with the AMS and the second four using the BESO algorithm. The Amin was set at a low value of 0.001mm2 for the AMS experiments while the initial mesh was populated with a diﬀerent number of elements for the four experiments. The four experiments using the BESO algorithm had diﬀerent ﬁxed meshes corresponding to the initial meshes of the AMS experiment. Optima the topologies for these experiments are shown in Fig. 5. Again, these topologies appear truss-like, noting that Exps. 11 and 12 correspond to Exps. 2 and 4 respectively.Table 2: Sensitivity test of AMS and BESO algorithms involving eight experiments (Exps. 5 to 12) characterized by their M , Dmin, Lmin, C∞ and K∞. Standard errors, σm, are also shown.Exp. Algorithm Start No. elems.

5 to 12) characterized by their M , Dmin, Lmin, C∞ and K∞. Standard errors, σm, are also shown.Exp. Algorithm Start No. elems. Amin5 6 7 89 10 11 12AMS AMS AMS AMSBESO BESO BESO BESO2,000 20,000 40,000 75,0002,000 20,000 40,000 75,000(mm2) 0.001 0.001 0.001 0.001 σm: - - - - σm:M (mm) 28 26 31 28 1.03 10 24 25 35 5.14Dmin Lmin (mm) (mm) 0.37 0.36 0.42 0.38 0.38 0.23 0.32 0.34 0.02 0.03 10.66 3.64 1.20 1.66 0.38 0.38 0.64 0.74 2.48 0.73C∞ (Nmm) 2.25 2.29 2.31 2.302.09 2.18 2.21 2.23K∞ (N/mm) 27.11 27.18 26.99 27.10 0.04 26.23 26.79 26.84 26.94 0.32It was diﬃcult to ascertain the sensitivity of the AMS and the BESO to initial mesh by simple vi- sual inspection of the topologies. Therefore, these topologies were quantiﬁed with the number of strut members, M , occurring in the optima (table 2), the minimum member thickness, Dmin, the minimum member length, Lmin and K∞.

Therefore, these topologies were quantiﬁed with the number of strut members, M , occurring in the optima (table 2), the minimum member thickness, Dmin, the minimum member length, Lmin and K∞. For both the AMS and BESO algorithm, variability in each sample set was quantiﬁed with the standard error in the mean, σm, of these characteristics stated as,σm =σ √ 4(12)(a)(b)(c)(d)(e)(f)(g)(h)Figure 5: Optima topologies for (a) Exp. 5 (b) Exp. 6 (c) Exp. 7 (d) Exp. 8 (e) Exp. 9 (f) Exp. 10 (g) Exp. 11 (h) Exp. 12where σ was the standard deviation of each sample set. The σm for both algorithms can be seen in table 2. These values show that there is less variability in the topological data collected for the AMS than with the BESO algorithm. Secondly, it was observed that the AMS realized members with smaller Dmin and Lmin than the BESO.

Secondly, it was observed that the AMS realized members with smaller Dmin and Lmin than the BESO. Attaining these dimensions with the BESO, or any other ﬁxed mesh algorithm, would require a much ﬁner mesh, and have high computational time. These ﬁner details are important as we are now able to produce part with such details via AM. A typical limit of minimum feature size in AM is about 10µm [32]. With regards to C∞, these could not be reliably used to estimate variability in the performance for the reasons mentioned in the previous section. A useful measure was the standard error in the K∞, which was also observed to be less for the AMS than that of the BESO as shown in table 2. It should also be noted that Figs. 3a and 3b, Figs. 3c and 3d shared similar boundary roughness since the size of boundary elements in both sets were largely similar. However, Figs. 3a-3d had smoother boundaries Figs.

3a and 3b, Figs. 3c and 3d shared similar boundary roughness since the size of boundary elements in both sets were largely similar. However, Figs. 3a-3d had smoother boundaries Figs. 3e-3h since elements at their boundaries were much smaller (0.001mm2) than that of the uniform meshes used in Figs. 3e-3h. This demonstrates that the AMS is able to achieve smoother boundaries than the BESO algorithm at a lower computation cost.Conclusion and further workA hybrid TO algorithm is proposed constituted by a modiﬁed form of the BESO algorithm and an adaptive meshing strategy (AMS). The BESO algorithm was modiﬁed so that it was based on strain energy densities since the AMS caused a progressive variation in elemental sizes. A two dimensional plate problem was solved to benchmark performance of the hybrid against that of the BESO algorithm.

A two dimensional plate problem was solved to benchmark performance of the hybrid against that of the BESO algorithm. The AMS was found to be substantially more eﬃcient, with a reduction in computation time of 52% and 82% respectively for the cases investigated, with similar performance. It was also less sensitive to the starting mesh and able to attain smoother boundaries with a lower number of elements. Also, ﬁner features are achievable with the AMS, starting with relatively coarse meshes. This is particularly interesting for topologies to be made via additive manufacture (AM). Since AM can achieve parts with high degrees of complexity. Further work could involve the determination of an element size that corresponds to AM’s resolution.

Since AM can achieve parts with high degrees of complexity. Further work could involve the determination of an element size that corresponds to AM’s resolution. Also, the AMS could be extended to practical three dimensional parts experiencing single and multiple load cases.AcknowledgementsThe authors are grateful for the funding provided by the IMCRC and EPSRC.References[1] Mitchell, A.G.M., The limits of economy in frame structures, Philo, Mag. Sec, 6(8), p.589-597 (1904).[2] Rozvany, G.I.N., Aims, scope, methods, history, and uniﬁed terminology of computer-aided topologicaloptimization in structural mechanics. Struct Multidisc Optim, 21, p. 19 (2001).[3] Kirsh, U., On optimal topologies of grillage structures, Engineering with computers, 42(8), p.223-239(1986).[4] Bendsoe, M.P., Kikuchi, Generating optimal topologies in structural design using a homogenizationmethod. Computer Methods in Applied Mechanics and Engineering, 71, p.

brid Automata in the Context of Additive ManufacturingJ.C. Boudreaux Guest Researcher National Institute of Standards and Technology U.S. Department of CommerceAbstract. To maintain the forward momentum of additive manufacturing technology, it is necessary to thoroughly evaluate new and potentially useful technological developments in this field. One such development is the intense interest being directed to the field of hybrid automata (HA). Hybrid automata combine both the discrete processing behavior of finite automata as well as the continuous, or flow, behavior of dynamical systems. At this point, some important results on hybrid automata have been obtained, but many open questions remain, including those concerning the decidability of HS operational procedures. (Recall that decidability is directed to a decision problem, that is, a definite true-or-false response given by an effective procedure.) Some important decidability results for HAs have been obtained.

For example, in [Henzinger et al.1998] the reachability problem for timed automata (an HA class) has been convincingly shown to be decidable. However, it should also be noted that subtle and difficult issues have been identified, e.g., [Fraenzle 1999], [Asarin, Collins, 2005]. This paper will provide a summary review of the operational features of HAs as they might pertain to additive manufacturing, and then briefly consider the following technical issues: (i) are the classical models of the real numbers best suited to deal with the necessarily approximate measures of physical systems or would non-standard analysis of [Robinson 1996] be a better fit; and (ii) would the introduction of “noisy semantics” and finite arithmetic precision, following [Freidlin, Wentzell 1984], be a better work around?I.

IntroductionAdditive Manufacturing (AM) has had significant impact on the production of engineeringprototypes and more recently on the manufacture of functional parts, especially the direct fabrication of molds, dies, and other tooling.

As now practiced, this style of manufacturing is based on additive manufacturing processes which may resolved into the following steps: first, to create a solid geometric model of the part to be built and to select a part orientation; second, to form an ordered sequence of planar slices which in effect decompose the solid into a sequence of thin cross-sectional polyhedral layers; and then to fabricate the part by producing the polyhedral designs by applying such methods as stereolithography, selective laser sintering, laminated object manufacturing, or fused metal/ceramic deposition [7], [8].Within the now-dominant AM paradigm, parametric design is based on the premise that aninventory of existing designs can be used to create design templates such that (1) the identified functional connections are coded by embedded parametric expressions, and (2) novel variant designs may be generated by assigning reasonable values to the parameters and then evaluating the resulting expressions in a precisely controlled manner.

This program has proven to be significantly more difficult than anyone had expected. This paper is an attempt to construct a qualitative parametricengineering model of the product design process in the context of discrete part manufacturing. It seems natural to suppose that engineered artifacts may be resolved into components parts that can be usefully represented as compact connected 3-manifolds with boundary where the boundary consists of a possibly disconnected set of compact smooth 2-manifolds (topological surfaces).

This style of manufacturing is based on additive manufacturing processes which may resolved into the following steps: create a solid geometric model of the part, defining surface boundaries and identifying interior (material) regions from the exterior (void) regions; select a part orientation and then form planar slices, decomposing the solid into a sequence of thin cross-sectional polyhedral layers; and fabricate the part by producing all of the polyhedra by any of several methods, including: stereolithography, selective laser sintering, laminated object manufacturing, fused deposition modeling, and three dimensional printing. In classical manufacturing work is applied to the surface of the part either through deformation process or by material removal. But in the present context the most useful characteristic of AM is that it reduces the 3-space shape of compact objects to a sequence of 2-space slices (and possibly some mild “skinning” conditions). [8]A.

[8]A. Geometric AM Models and Applied Topology.But there is still more to be observed here: the stack of planar slices, properly understood, is ofvery great importance in its own right. Not only does it allow manufacturers to produce that one specific part, but also, suitably transformed, it has the potential to produce infinitely many geometrically distinct parts, every one of which is topologically identical to the single specific part with which we began. This important notion should be more precisely defined. [19]A topological space is a pair (S, O) where S is an unempty set of points and O is a set of subsetsof S, called the open sets, which satisfies the following conditions: O contains both S and the empty set , and is closed under finite intersections and arbitrary unions. A closed set is the set-theoretical complement of an open set.

A closed set is the set-theoretical complement of an open set. The collection C of all closed sets satisfies the dual conditions: C contains both  and S, and is closed under finite unions and arbitrary intersections. Let Y be an arbitrary subset of S and p be a point, then p is a limit point of Y only if every open set of (S, O) contains a point of Y distinct from p. That is, elements of Y get arbitrarily close to p. Now suppose that S and T are topological spaces, then a function f from S to T is said to be continuous, or a map, provided it is an into function and for every subset X of S and point p, if p is a limit point of X, then f(p) is a limit point of f(X), the image of X under f. Equivalently, f is continuous if it is into and the inverse image under f of every open set of T is an open set of S.

Equivalently, f is continuous if it is into and the inverse image under f of every open set of T is an open set of S. Given this preliminary discussion, the fundamental meaning of topological identity is finally made clear: a function from spaces S to T is a homeomorphism provided it is a one-to-one onto function (a bijection) and both the function and its inverse are continuous.Then, with the topological sketch in place, we can fashion at will an infinite variety oftopologically identical geometric constructions in two steps: first, we can arbitrarily thicken all (or some) of the initial planar slices, while preserving all structural elements (such as holes or voids); and second, without tearing or puncturing through the thickened slices, we can reshape them at will, just as we would reshape wet clay … well almost!B.

Resemblances (1): Morse Theory.We are not quite finished yet! What needs to be added to this story is some structuring principle which allows all of the infinitely many possible geometric constructions to at least resemble one another. Suppose that the geometry of our first part was a child’s doll, say Raggedy Ann, then what we would like to have is a mechanism in place which would ensure that resemblance to Raggedy Ann would be at least roughly approximated. One way to do this is to fix the doll in a plausible fixed position, and then to scan her from just above her head (the initial scan point) to just below her shoes (the final scan point) while noting for each scan the distance from the initial scan. Of course, the results would be an ordered set (top-down) of planar slices! [6]Again, we’ll need a bit more topology [6], [7], [19].

Of course, the results would be an ordered set (top-down) of planar slices! [6]Again, we’ll need a bit more topology [6], [7], [19]. An n-dimensional manifold (n-manifold) Mis a separable Hausdorff topological space such that for every point p, there is a neighborhood of p, which is homeomorphic to Euclidean n-space or to Euclidean n-halfspace. Let U be an open set for some point p of M and let h be a homeomorphism from U into n-space, or equivalently, any open n- ball, then the pair (U, h) is called a chart of dimension n. An atlas A on M is a collection of charts whose open sets cover M. Suppose that (U, h) and (V, k) are charts whose open sets have a nonempty intersection, say W, then h and k given rise to a homeomorphism on the n-space images of W , specifically, (kh -1): h(W) k(W), and which establishes functional relationships between the n-tuples in the respective images of W.

If all of the functions so established are differentiable (in class Ck or C) then M is said to be a differentiable (smooth) manifold.The collection of points (if any) with neighborhoods homeomorphic to the n-halfspace definethe boundary of the manifold, M. Then given two topological spaces X and Y, let f,g: X  Y be continuous functions. Then f and g are homotopic iff there is a continuous mapping h: X  [0,1]  Y such that for all x in X, h(x,0) = f(x) and h(x,1) = g(x). The mapping h is a homotopy between f and g and the product space X  [0,1] is the homotopy cylinder.

The mapping h is a homotopy between f and g and the product space X  [0,1] is the homotopy cylinder. The existence of a homotopy between functions establishes that the functions can be continuously deformed into one another.The final piece that needs to be added to this story is that there needs to be a closer connectionbetween the topological structure of the compositions that we’re proposing to manufacture and the critical points of smooth functions on the manifold (points at which all partial derivatives of the function are null). Critical points consist of maxima, minima, and saddle points. First, assume that M is a nice 3-mainfold and that M, the boundary of M, is the disconnected union (possibly empty) of nice manifolds V and W. Intuitively, think of V and W as planes which sandwich M between them. Then let f be a smooth mapping of M into a real closed interval [a,b] such that f -1(a) = V and f -1(b) = W.

Intuitively, think of V and W as planes which sandwich M between them. Then let f be a smooth mapping of M into a real closed interval [a,b] such that f -1(a) = V and f -1(b) = W. Second, generate slices through M, that is, if x in [a,b], then f -1(x) is such slice. Not all such functions are useful. What we want to do is to restrict our attention to Morse functions whose critical points on M are non-degenerate (the Hessian matrix of partial second derivatives is non-singular at the point). On nice manifolds, these functions are plentiful. [6], [19]The point of this exercise is to use f to scan M to build a CW-complex that has the samehomotopy type as M. The cell complex is built up by a sequence of cell attachments, which is completely determined by the sequence of critical points. As f slides between critical points, the topology of the manifold does not change, that is, no cells need to be attached.

As f slides between critical points, the topology of the manifold does not change, that is, no cells need to be attached. When f reaches a critical point, the topology of the manifold changes in way that requires the cell complex to be updatedby attaching a k-cell, where k is the index of the critical point (the number of negative eigenvalues of the Hessian matrix). Since nice manifolds are compact, there are at most finitely many critical points, which implies that the sequence of critical points pi and the associated critical slices f -1(xi) completely encode the intrinsic topology of M.C. Resemblances (2): Reeb Graphs.Once a Morse function f has been defined on M, a Reeb graph can be used to generate a skeleton of M [6], [19]. A quotient space is formed on M  R by identifying points, which belong to the same connected component of a slice.

A quotient space is formed on M  R by identifying points, which belong to the same connected component of a slice. That is, if p and q points, then (p, f(p)) is identified with (q, f(q)) if f(p) = (f(q) and p and q are in the same connected component of f -1(f(q)). The nodes of the Reeb graph will be the critical points of M and the edges will be given by the connected components between critical slices. From the perspective of Morse theory, the most interesting slicings are those which (1) begin and end with minimal patches and which (2) mark the critical regions of the 3- manifold under study, that is, regions in which either the number of components or their topological properties change. An example of a change of the second kind is the first appearance of a new boundary patch (or void surrounded by boundary elements) in a patch whose earlier stages had none. Let slicings of this kind be called Morse slicings. In general, 3-manifolds admit of many Morse slicings.Fig.

Let slicings of this kind be called Morse slicings. In general, 3-manifolds admit of many Morse slicings.Fig. 1 : Reeb graph of torus Fig. 2 : Reeb graph of 2-torusIn effect, what Reeb graphs really do is to create stratified structures, that is, a singlecomponent at the first (or top) level, which is called the root of the structure and which may be refined into a branching networks of components (the descendants) at the lower levels. The condition which must be satisfied is that every descendant node must be linked to the parent node by a unique sequence of links. Such organization structures are called rooted trees. Stratification also leaves the parent’s boundary intact, allowing it to become a container of the boundaries the children.Boundaries drawn in this manner impose a nest structure on the product: at the highest level isthe product itself, followed by its immediate descendants, and so on, until all of the boundaries have been accounted for.

In the artifact-as-network model, stratification boundaries mark components, which can then be identified by symbolic “tags” or labels. This symbolic structure permits stratified product networks to form a framework with respect to which the design features of components are defined. Features of one component may be either synthesized from those of its descendant components in a defined stratification or inherited by virtue of its links to other (in some cases, all other) components.Given a 3-manifold K and a Morse slicing (X1,...,Xm), it is possible to construct a combinatorial variant of a Reeb graph.

Under the assumption that the slicing has already been reduced by eliminating consecutive pairs of homeomorphic slices, then(1) the vertices of the Reeb graph are components of slices; and (2) the arrows of the Reeb graph link each component of slice Xi to all of the components of slice Xi+1 which descend from it.As in the smooth case, Reeb graphs concisely encode important topological properties of 3-manifolds: the vertices encode the components of slices, and the arrows encode the 3-manifolds (like cylindrical solids) which link a component of slice Xi with all of the components of slice Xi+1 from which descend from it. The 3-manifolds paired with Reeb graph arrows may have voids and tunnels. However, Reeb graphs are less successful in capturing properties affecting the geometric realizations of 3-manifolds.For example, there are many ways to embed an encoded Reeb graph into 3-space which are notgeometrically equivalent (isotopic) to one another.

This has to do with the mathematics of knots in 3- space. The multiplicity of these relational ties goes a long way in explaining the observed conceptual diversity among feature categories.Again, functional features are inherited from above by considering the constraints imposed bythe component's children. Specifically, the functional capabilities and operational behavior of components may be modeled by transfer functions which, given time-sequences of inputs (and possibly information about the current state of the component), produce time-sequences of outputs (and possibly a new state) after a specified time lag. Mating features are both synthesized from below and inherited from above because of the requirements imposed by assembly and joining processes. Form features are those through which all of the (often) contending requirements of the part are finally reconciled and by which the connectivity-induced constraints are resolved.II.

Form features are those through which all of the (often) contending requirements of the part are finally reconciled and by which the connectivity-induced constraints are resolved.II. On the Architecture of AM: Two Recent ReportsThe “planar slice” AM architecture remains a central paradigm and undoubtedly will continueto develop as it has done now for a quarter of a century. A very interesting example of current developments AM was described in a recent report in Science. Zheng et al. [21] have reported a wayto build ultralight and ultrastiff mechanical materials by constructing a class of cellular materials, thereby mimicking natural foamlike structures, such as trabecular bone, plant parenchyma, and sponge, each of which combines low weight with superior mechanical properties.

By reducing the degree of random porosity occurring in natural materials, they proposed to develop a class of materials that contains micro- and nanoscale building blocks arranged in an ordered hierarchy. The authors report that there were several proposals for the unit cells, including an octet-truss (a stretch-dominated structure) and a tetrakaidecahedron (a bend-dominated structure). The critical feature sizes of the unit cells ranged from 20m to 40nm. The fabrication of the microlattices was enabled by projection microstereolithography, a layer-by-layer AM process which they describe as able to build “arbitrary three dimensional structures.” The authors report that the specific stiffness of the Ni-P stretch- dominated metallic lattice stays very nearly constant, measured at 1.8  106 m2 /s2 and 2.1  106 m2 /s2 at densities of 14 mg/cm3 and 40 mg/cm3, respectively.

The article concludes with the claim that “fabricating ordered lattice structures at these lengths scales brings them into the regime in which it becomes possible to design microstuctured functional materials with superior bulk-scale properties.”This interesting report is one example among many of the vigor of the AM engineering community. In thinking about this more carefully, and especially by happening across the Biomedical Engineering Society (BMES), it struck me, as it has obviously struck many others, that a strongly supported collaboration between BMES (and similar groups) and the AM community would be mutually beneficial. The 2014 paper by Kamm and Bashir [16], entitled “Creating Living Cellular Machines,” begins with the following comments: “ One of the greatest opportunities lies in the potential to understand populations of multiple cell types and their interactions.

… Yet, it can be argued that a tremendous gap exists between understanding processes at the level of a single cell and the behavior of large-scale tissues, i.e. how he local rules of interaction result in global functionalities and diverse prototypes. This is an issue involving complex systems of multiple interacting components that could fruitfully draw upon considerable advances in the engineering realms of forward design and manufacturing of large, complex systems.” [16, p445] This paper suggests that a significant direction of the future growth for AM technology could be enhanced by investigating “soft” architectures, which are likely to dominate the biological approaches. In applications of this sort the analogue of the granular “planar slice” would be “soft” polymers, hydrogels, or other biomaterials.

In applications of this sort the analogue of the granular “planar slice” would be “soft” polymers, hydrogels, or other biomaterials. I presume that the introduction living cellular material would be accomplished by an injection process, and that this could result in local but significant deformation of the “soft” material. In this case, it also seems that mathematically useful description of the process is one which allows a deformable topological process, perhaps, a structural approach that jointly handles both the temporal and the metric properties of the space. This task is made no easier when such added features as cell motility and intercellular chemical diffusion between the introduced samples are taken into account.III. Hybrid AutomataA.

This task is made no easier when such added features as cell motility and intercellular chemical diffusion between the introduced samples are taken into account.III. Hybrid AutomataA. Toward a Definition of Hybrid Automata.In section IV below, I intend to sketch out the view that this interesting class of hybrid automata, whose defining characteristic is that they are designed to function in both discrete and also in continuous modes, are plausible candidates for the role of supervisory controllers in the context ofadditive manufacturing. I put this proposal forward tentatively in the hope that the technical issues which would need to be resolve to accomplish this goal are still under very active discussion. [4], [13], [15]The general notion of hybrid automata (HAs)is that they are able to display both discrete (orjump) as well as continuous (or flow) dynamics.

[4], [13], [15]The general notion of hybrid automata (HAs)is that they are able to display both discrete (orjump) as well as continuous (or flow) dynamics. The technical definition of hybrid automata begins by introducing directed graphs, usually abbreviated to digraphs. More precisely, a digraph is a structure (V, E) in which V an unempty set of vertices, which are also called nodes, and E is a possibly empty set of edges. It is also understood that the vertex set and the edge set of a digraph are mutually disjoint. It will be assumed that the edges have a definite direction, that is, every edge points from one vertex to another. If v and w are connected vertices, then edges between them may have one of two possible directions: v  w or v  w, and this means that every edge is oriented. [2], [3], [5], [9]The definition of digraphs given above admits a very large class of structures in terms of whichthe inter-component relational ties can be explicitly accounted for.

[2], [3], [5], [9]The definition of digraphs given above admits a very large class of structures in terms of whichthe inter-component relational ties can be explicitly accounted for. For example, digraphs may have many (even infinitely many) distinct parallel and antiparallel edges between the same two vertices. The first step is straightforward: the vertices corresponding to assembled components are symbolically “tagged” and the arrows consist of both organization links that connect parents with their children and also a complex web of links between components.To properly understand the implications of this 2-way hybrid structure, note that each vertex can be interpreted both as a place in which one or more actions are initiated by the node's immediate predecessors (if any). The interweaving of product and process data in a single structure is very useful. In effect, it supplies a single unified support system for many of the operational control methods.

The interweaving of product and process data in a single structure is very useful. In effect, it supplies a single unified support system for many of the operational control methods. For example, if the digraph is interpreted in terms of the processes to be performed, then it contains the information needed to identify and initiate the actions that will be needed to produce the required effects. Hybrid automata also require that there may be a multiplicity of edges between any two vertices, that is, the digraphs which define that basic groundwork of hybrid automata are multidigraphs. The only edge configurations which are explicitly disallowed are loops, that is, edges which are directed from one vertex to itself.

We now have enough basic information to proceed to a suitable definition of the functional components of formal model of computation.(HA1) Control Digraph: a finite directed multidigraph (V, E) whose vertices are called control modes, or locations, and whose edges are called control switches.(HA2) Variables: (i) a finite set X = {x1 , x2 , … , xn } of real-valued variables, where the index n is the dimension of the hybrid automata; (ii) a finite set DX = {Dx1 , Dx2 , … , Dxn }, which are first derivatives of the variables (with respect to time) during continuous change; and (iii) a finite set +X = { +x1 , +x2 , … , +xn } of “+” marked variables which represent values at the conclusion of discrete change.(HA3) Initial, Invariant, and Flow Conditions: Three vertex labeling functions init, inv, and flow that assign to each control mode (i.e., to each vertex v in the control digraph) three predicates: (i) init(v) is a predicate whose free variables are in X which assigns initial conditions to each control mode; (ii) inv(v) is a predicate whose free variables are in X which assigns invariant conditions to each control mode; and (iii) flow(v) is a predicate whose free variables are in X  DX which assigns invariant conditions to each control mode.The rule (HA3) explicitly introduces the logical notion “predicate” which is properly understood as an expression which is applied to one or more arguments, i.e., either variables that may be assigned different values from one time to another or constants whose values are fixed in advance.

Unlike functions, the values returned by the application of predicates are limited to a specific range of truth- values, usually limited to “true” or “false” but which can be extended to include a much larger collection of allowable values.(HA4) Events: a finite set Σ of events and an edge labeling function event: EΣ that assigns a suitable control switch (i.e., to edges in the control digraph) an event.(HA5) Jump Conditions: a edge-labeling function jump that assigns to each control switch e in E a predicate jump(e) whose free variables are in X  +X.B. Putting Hybrid Automata to Work.Having identified the components, the next step is to describe execution of hybrid automata.The first, and most obvious point, is that there are two operational modalities: a continuous modality in which the process is a flow, which may be pictured as a kind of evolution, and a modality in which the change is a discrete jump.

[2], [3], [9], [11], [14], [15]The idea is that both modalities can be properly handled by accepting variant of the familiarclass of labeled transition system as an operational paradigm. In this case, we first identify a state space Q, which may be infinite, and a subset Q0 of all initial states, then identify a (possibly infinite) set Rel of binary relations rj  QQ. Each binary relation is a set of ordered pairs of states, so for any rj we may distinguish dom(rj), the set of all first members of the relation, and rng(rj), the set of all second members of the relation.The operation of a labeled transition system is not straightforward: (i) pick any initial state in Q0, (ii) given that state, say q0, select from Rel any relation rj such that q0 is a element of dom(rj) and then add to the sequence of states the corresponding element of rng(rj); (iii) continue (perhaps to infinity) until no further additions are possible.

This procedure is only roughly described in this case and a bunch of qualifications and distinctions would need to be made.The useful observation is that there are two sequences involved: the sequence of states from Q and also the sequence of relations from Rel. Having sketched labeled transition systems in general, the next task is to focus more narrowly on hybrid systems and in particular on timed transition systems for n-dimensional hybrid automata of the form (Q, Q0, A, Rel):(HA6) The state space of an n-dimensional hybrid automaton is defined as Q, Q0  V  Rn, where Rn is an n-dimensional vector space. Thus, the elements of Q and Q0 are ordered pairs of the form (v, x), where v is a control mode (i.e., a vertex) and x is an n-dimensional vector.(HA7) The edge label set of an n-dimensional hybrid automaton is defined as A  Σ  R0, that is, the union of all of the edge labels defined in (HS4) and the set of all non-negative real numbers.

(Recall that the set components of A are taken to be mutually disjoint.)(HA8) (i) The ordered pair (v, x) is an element of Q iff the closed predicate init(v)[X = x] is true; and (ii) the ordered pair (v, x) is an element of Q0 iff both init(v)[X = x] and inv(v)[X = x] are true.(HA9) For each event  in Σ and for each rk in Rel, rk((v, x), (v’, x’)) iff there is a control switch e in E such that (i) the edge e is from v to v’, (ii) the closed predicate jump(e)[ X, +X = x, x’] is true; and (iii) event(e) = .(HA10) init and inv Rules: (i) The ordered pair (v, x) is an element of Q iff the closed predicate init(v)[X = x] is true; and (ii) the ordered pair (v, x) is an element of Q0 iff both init(v)[X = x] and inv(v)[X = x] are true.(HA11) jump (Discrete) Transitions: For each event  in Σ and for each rk in Rel, rk((v, x), (v’, x’)) iff there is a control switch e in E such that (i) the edge e is from v to v’, (ii) the closed predicate jump(e)[ X, +X = x, x’] is true; and (iii) event(e) = .(HA12) flow (Continuous) Transitions: For each nonnegative real  and for each r in Rel , r((v, x), (v’, x’)) iff v = v’ and there is a differentiable function f : [0, ]  Rn , with the first derivative Df : (0, )  Rn , such that: (i) f(0) = x and f() = x’ ; (ii) for all reals  in (0, ), both init(v)[X = f()] and flow(v)[X, DX = f(), Df()] are true; and (iii) the function f is called a witness for the transition r((v, x), (v’, x’)).IV.

A Model for Supervisory Hybrid AutomataThe discussion in the previous section has sketched in outline the basic operational structure ofhybrid automata. The main effort in this section is going to be devoted to a consideration of the possible application of hybrid automata theory to the supervisory control of continuous systems. This requires reasonably well-defined linkage between a continuous plant on one end and a discrete system on the other. One obvious problem is that the ends of this linkage do not speak the same language! That is, the behavior of the plant is governed by a well-defined family of ordinary differential equations, but the computational end of this configuration, usually taken to be a discrete event system (DES) which is in fact an event-driven finite-state automaton.

[18]Therefore, in order to communicate at all, there needs to be a bidirectional interface whichmoves (1) the analog-to-digital signals (i.e., the continuous-time output or plant-state) from the plant to the DES, and also (2) the digital-to-analog signals (i.e., the discrete-time output, or DES-state) from the DES to the plant. Having this distinction available is useful, but there are still fundamental differences between the continuous realm of plant-states on the one hand and DES-states on the other.Let’s first approach this from the perspective of the plant, and specifically the notion of a plant-event. As explained in [18], “a plant event is simply and occurrence in the plant… In the case of hybrid control, a plant event is defined by specifying a hypersurface that separates the plant’s state space into two disjoint sets.

The plant event occurs whenever the plant state trajectory crosses this hypersurface.”Though not explicitly introduced here, the inherent structure of plant space is an n-dimensional vector space through which the plant-event moves and the overarching mathematical framework is that of dynamical system theory. The authors of [18] also make it quite clear that “crossing the hypersurface” is not a simple matter to define in practice: (1) the trajectory could have a very brief excursion across the hypersurface and could even repeatedly re-cross it, and (2) the normal practice would require an unambiguously clear crossing event. The authors propose that in order for a crossing to be recognized, a set of smooth functional should be used to disambiguate the situation ([18], 1029).

The authors propose that in order for a crossing to be recognized, a set of smooth functional should be used to disambiguate the situation ([18], 1029). If all has gone to plan, then the bidirectional interface uses the generator subsystem to convert the continuous-time plant-state to an asynchronous symbolic input to the DES control system.From the perspective of the DES, the operation of the bidirectional interface accepts symbolic inputs from the actuator subsystem and then produces appropriate symbolic outputs for the generator. It should be noted that behavior of the DES system is a matter of a explicit computational conversion from input symbolic-text (for the actuator) to output symbolic-text (for the generator).At this point, the notion of a system controller, perhaps better called the supervisor, can bestraightforwardly introduced.

Here is the main line taken: there is in the world a physical plant whose operational style depends upon life in a bounded continuum, and also in the world there is a supervisor whose operational style depends upon life in symbolic realm of at most denumerably infinite symbolic structures, and finally there is in the world a supervisor whose operational style is to shuttle symbolic expressions from one part of itself (the actuator part) to another part of itself (the generator part)subject to the condition that this shuttling shall always and everywhere carry finite and only finite messages.V. DiscussionFirst, the easiest generalization about AM manufacturing is that, if and when anything getsmade, it will almost inevitably get made in a closed box.

DiscussionFirst, the easiest generalization about AM manufacturing is that, if and when anything getsmade, it will almost inevitably get made in a closed box. And the second easiest generalization about AM manufacturing is that any one closed box will have virtually no contact with whatever happens in any other closed box.I do accept the point that the discussion of supervisory HAs, particularly in Section III.C, was abit of a stretch, but in a larger sense what I am recommending is that AM technology be fully incorporated into the industrial engineering environment.The commercial viability of AM and the ability of this technology to add value to such productclasses as automobiles, household appliances and industrial machine tools is now widely accepted. Functional requirements and detailed specifications need to be developed.

Functional requirements and detailed specifications need to be developed. Validation procedures will be needed to assure that the functional requirements are satisfied by interlinked networks of HAs.AM technology can offer more cost-effective and higher-quality solutions compared totraditional devices with similar functionality, or they offer possibilities that cannot be realized at all using traditional systems. Compared to the traditional devices, AM can be produced from small amounts of raw materials, use little energy and generate small volumes of waste products. But it must also be clearly understood that there are substantial technical and engineering barriers to be overcome.Second, this paper, more perhaps than earlier ones, has brought home to me the need to takegreat care in the construction of important mathematical notions.

I am now focusing on Non-Standard Analysis (NSA) – and I believe that this notion will be an increasingly powerful tool in the engineering tool box. This work was developed by several authors including Skolem and also Laugwitz among others, but the main systematic development of NSA was achieved by Robinson, esp. in his book Non- Standard Analysis, which was published in 1966. Here’s a very brief summary of this mathematical idea (taken from Alberverio et al.[1]): the main result of NSA is that “the geometric line or continuum can support a point set richer than the standard reals. This … gives us a framework for a geometric analysis of physical phenomena on many scales …”This book was well-received, especially by those mathematicians who had been acquainted with Model Theory. Those less familiar with this logical material were put off by such notions as ultrafilters, ultraproducts, and other seemingly alien notions.

Those less familiar with this logical material were put off by such notions as ultrafilters, ultraproducts, and other seemingly alien notions. In the early days it was commonly believed that anything that could be done within NSA could also be done in analysis without using NSA techniques. However, it soon became obvious that this idea (or hope) was not true. Today NSA is showing signs of robust health, especially in the area of applied analysis and related areas of mathematics which have direct and immediate bearing on the engineering community.There are other areas of mathematics which should also be mentioned.

One very interesting example is the work of Stefan Hilger, who developed the notion of measure chains (also called time scales) in order to be able to unify the study of differential equations and difference equations into a single theory.Finally, I am fortunate to have been given the opportunity in my present role at NIST to be able to explore these intriguing areas more thoroughly and more deeply than would otherwise have been possible.References[1] S. Alberverio, J.E. Fenstad, R. Hoegh-Krohn, and T. Lindstrom, Nonstandard Methods in Stochastic Analysis and Mathematical Physics, Academic Press, 1986.[2] R. Alur, et al. “The Algorithmic Analysis of Hybrid Systems,” Theoretical Computer Science, vol 138 (1995); 3-34.[3] R, Alur, et al., “Discrete Abstractions of Hybrid Systems,” Proc. IEEE, vol 88 (2000); 971-984.[4] E. Asarin and P. Collins, “Noisy Turing Machines,” in L.

von Neumann, “Probabilistic logics and the synthesis of reliable organisms from unreliable components,” in C.E. Shannon and J. McCarthy (eds.), Automata Studies, Princeton, 1956; 329-378.[21] X. Zheng et al., “Ultralight, Ultrastiff Mechanical Metamaterials,” Science, 20 June 2014, vol(334) 2014; 1373-1377.Biography. Dr. Boudreaux joined the National Institute of Standards and Technology, formerly the National Bureau of Standards, in 1979, first in the Computer Science Laboratory, and then from 1984 to 1992 in the Manufacturing Engineering Laboratory. In 1992, he was appointed a program manager in the NIST Advanced Technology Program. His research has focused primarily in the following areas: models for higher-order functional systems, adaptive control, real-time error compensation systems, and intelligent manufacturing systems. Contact information: J.C.

viewed PaperHYBRID CURVE FITTING FOR REDUCING MOTION COMMANDS IN OBJECT CONSTRUCTIONCharles Wade* and Michael Borish* *Manufacturing Science Division, Oak Ridge National Laboratory, Knoxville, TN 37932AbstractExisting slicing software for additive manufacturing typically requires a triangulated mesh as its input. Triangulated meshes are approximate representations of exact CAD models. Despite the loss in dimensional accuracy, triangulated meshes are used because they are computationally easier to cross-section and offset than the exact geometry in CAD format. When a triangulated object is prepared, the resulting machine instructions include only linear motion commands. Numerous modern motion controllers can move in arc and spline motions; however, the absence of slicing software that supports curvature prevents these commands from being leveraged.

Numerous modern motion controllers can move in arc and spline motions; however, the absence of slicing software that supports curvature prevents these commands from being leveraged. To address this limitation, this paper presents a method for the hybrid reconstruction of arcs and splines as a post-processing step to traditional slicing. This method can greatly reduce the number of motion commands required to construct an object by printing smooth curved surfaces. This concise representation of tool-pathing allows for more even extrusion and is computed without a major impact on slicing time.IntroductionObjects in additive manufacturing (AM) are normally constructed using sequences of line segments that define the geometry of the object. Toolpaths are commonly exported by slicing software as g-code instructions G0 and G1 that are interpreted as linear motion commands by a machine.

Toolpaths are commonly exported by slicing software as g-code instructions G0 and G1 that are interpreted as linear motion commands by a machine. As demonstrated in Figure 1, line segments are generated as part of the cross- sectioning process that occurs to cut 3D objects into toolpath layers [14]. Since the operations to cross-section triangulated meshes and offset polylines are simpler than the corresponding curved geometry operations, most slicing software uses an exclusively line segment-based approach. To accurately represent curved geometry, these segments use a high density of linear commands to approximate arcs and splines. The sheer number of line segments required for objects with curved surfaces results in sizeable g-code files that become infeasible for large format AM control systems. Additionally, densely packed line segments result in choppy machine kinematics that reduce both the quality and probability of a successful print.

Additionally, densely packed line segments result in choppy machine kinematics that reduce both the quality and probability of a successful print. Control systems that are computationally limited may also be incapable of processing successive straight segments at high speed, increasing construction time. Furthermore, approximating curved geometry with line segments reduces the overall surface finish and tolerance of printed objects.In this paper, we present a hybrid approach for AM toolpath planning that leverages both arc and spline commands. The goal of this work is to identify sections in AM parts that contain curvature and apply fitting methods to replace densely packed line segments with arc and spline segments. First, we examine related and foundational work. Then, we provide an overview of the individual methods used for fitting arcs and splines.

First, we examine related and foundational work. Then, we provide an overview of the individual methods used for fitting arcs and splines. Next, we present a novel hybrid approach to determine when arc versus spline fitting should be applied to a section of line segments. Finally, we compare several examples constructed with only line segments and their corresponding hybrid-fit constructions. An analysis into the impact on slicing time is also conducted to determine the feasibility of integrating this algorithm into the AM workflow.This manuscript has been authored by UT-Battelle, LLC, under contract DE-AC05-00OR22725 with the US Department of Energy (DOE). The US government retains and the publisher, by accepting the article for publication, acknowledges that the US government retains a nonexclusive, paid-up, irrevocable, worldwide license to publish or reproduce the published form of this manuscript, or allow others to do so, for US government purposes.

DOE will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).680(a) (b)Figure 1: (a) Input triangulated mesh with cross-sectioning plane, (b) Cross-sectional layerapproximating curvature with many line segmentsRelated WorkTraditional AM slicing uses explicitly defined triangulated surface-mesh representations for 3D geometry. Although this representation reduces computational complexity with cross- sectioning, offsetting, and filling operations, it introduces numerous inaccuracies and space limitations. The use of implicit representations in AM is the subject of numerous papers [4, 5, 18, 19]. These works present interesting solutions for slicing with implicit functions, but none present a complete set of replacement operations for traditional AM slicing. For instance, Jin et al.

These works present interesting solutions for slicing with implicit functions, but none present a complete set of replacement operations for traditional AM slicing. For instance, Jin et al. proposed using the implicit level-set method to smooth input contours with curvature information [6]. Although the level-set method reduced the number of empty voids when compared with explicit slicing, it was limited to certain geometries and by computational complexity. Another approach was taken by Brunton et al. using an implicit model with signed distance fields [2]. Their algorithm primarily focused on space efficient representations of the input model for slicing, and approximating curvature using voxels. Despite providing good curvature approximation for systems well suited for voxel space like polyjet printing, their algorithm does not provide an efficient way to reduce line segments in traditional toolpath driven processes. Q. Li et al.

Q. Li et al. expressed the infancy of the implicit slicing topic in their review [7]. They found that most existing CAD tools are focused on subtractive manufacturing and lack support for 3D capabilities. For instance, most existing implicit subtractive tools are concerned with boundary representations. This is not well suited for AM processes that use various densifying structures to reduce material use or improve strength. While a completely implicit slicer replacement is widely sought after, there are many challenges that need addressed before one will become widely available.Given the current limitations of implicit slicing, there has been extensive investigation into fitting parametric curves to line segments [10, 12, 21, 23]. However, these papers have focused on traditional CNC path planning and there has been little work applying the techniques to AM.

However, these papers have focused on traditional CNC path planning and there has been little work applying the techniques to AM. Valdivieso proposed a greedy algorithm to collect possible arc candidate locations and fit them to681line segments in AM [16]. Leveraging the G2 and G3 commands supported by most AM motion controllers, Valdivieso’s work was primarily concerned with using biarcs to approximate curvature in an additively manufactured object. Most existing CNC toolpath planners rely on biarcs to approximate b-splines because they often lack support for the G5 command [3, 17, 22, 24]. Although this results in a reduction of motion commands over exclusively using line segments, it requires many arcs to represent dynamic curves. In contrast, graphical methods for curve fitting discrete data routinely employ splines [11, 15, 20]. Lin and Ding proposed a method to fit b-splines to ordered data points [9]. Similarly, W. Li et al.

Lin and Ding proposed a method to fit b-splines to ordered data points [9]. Similarly, W. Li et al. [8] presented an adaptive knot placement algorithm for b-spline curve approximation of discrete data.Consequently, this paper focuses on providing implicit capabilities through post-process curve fitting. This approach leaves the traditional slicing process intact, allowing the user to leverage standard toolpath planning algorithms for cross-sectioning, infill generation, and support structure calculation. The proposed method takes a series of line segment commands as input and returns a combination of arcs, splines, and lines that form an implicit representation of the toolpath. Input toolpaths can be either closed or open polyline contours but must contain at least two segments to fit a curve.

Input toolpaths can be either closed or open polyline contours but must contain at least two segments to fit a curve. A post-processing approach provides a non-intrusive method to leverage the toolpath planning capabilities in explicit geometry slicers, while enabling implicit construction with arcs and splines.Parametric Curve FittingArcsThe G2 and G3 commands define arcs using three parameters: a start point, an end point, and a center point. G2 is used for clockwise motion, while G3 is used for counterclockwise motion. When fitting an arc to a pair of consecutive line segments, the center point of the arc that circumscribes the segments must be determined. Figure 2 illustrates how the center point can be computed from two segments by finding the intersection of their perpendicular bisectors.Figure 2: Finding the center point of an arc fits segments Si and Si+1682Given a series of N ordered line segments, it can be determined if their radii of curvature match.

It is observed that consecutive line segments with identical curvature can be fit with a single arc. This is accomplished by picking a median point from the series and reducing the problem to a two-segment system. The center point can then be calculated for all points in the series. By partitioning a series of line segments into sections with matching curvature multiple arcs and lines can be fit optimally.B-SplinesThe G5 command uses four parameters to draw a b-spline: a start point, an end point, and two control points [13]. Although biarcs can approximate advanced curvature such as b-splines, they are less efficient than using the G5 command. Figure 3 shows that multiple G2 and G3 arc commands must be used to express the same corresponding G5 b-spline. The topic of fitting b- splines to discrete data points is the focus of multiple papers [8, 9]. To simplify this work and to focus on hybridized arc and spline fitting, the W. Li et al.

The topic of fitting b- splines to discrete data points is the focus of multiple papers [8, 9]. To simplify this work and to focus on hybridized arc and spline fitting, the W. Li et al. approach is used to determine the placement of b-splines for a series of line segments [8]. The algorithm proposed by W. Li et al. takes an ordered series of points, representing the start and end points of the line segments, and returns the optimal knot placement to construct b-splines.(a)(b) (c)Figure 3: (a) series of line segments, (b) fit with a single G5 b-spline, (c) fit with one G2 and one G3 arc683Hybrid FittingAlthough certain geometries drawn through the G5 b-spline command are more concise, G5 is not a universal replacement for constant radius G2/G3 arcs. In the trivial example of Figure 4, the circle can be drawn with a single arc command, where two spline commands to achieve the same result are required.

In the trivial example of Figure 4, the circle can be drawn with a single arc command, where two spline commands to achieve the same result are required. This exemplifies a need for a hybridized approach when fitting curves to line segments. The method for hybrid fitting first performs preprocessing to clean the input geometry. Next, it conducts initial partitioning based on an angle cutoff to preserve sharp contours. Then, curvature analysis is conducted to categorizes partitions based on their type. Finally, these sections are fit with the outlined arc and spline methods according to their category.(a)(b) (c)Figure 4: (a) Circular line segment path, (b) fit with a single arc, (c) fit with two splines684PreprocessingBefore a series of line segments can be fit with arcs and splines, it must be preprocessed to remove noisy and redundant data. Collinear segments do not contribute toward the hybrid fitting process and must be replaced with a single line segment.

Collinear segments do not contribute toward the hybrid fitting process and must be replaced with a single line segment. Likewise, segments of a very small length should be stitched with longer ones since they do not contribute substantially to the geometry. A linear smoothing parameter 𝛼 = 10 microns is used to identify sections of negligible length. However, this parameter depends on the scale of the machine. This work assumes that an input triangulated mesh was exported from CAD software with a reasonably high resolution. This assumption excludes models that were reconstructed from 3D scan or LIDAR data.Initial PartitioningThe initial partitioning phase identifies subsets of line segments that may be replaced with parametric curves. First, pairs of line segments that form sharp angles are detected. If fit with a curve, these pairs would result in over-smoothing of the path.

First, pairs of line segments that form sharp angles are detected. If fit with a curve, these pairs would result in over-smoothing of the path. Given the ordered list of segments, {𝑆𝑖, (𝑖 = 0, … , 𝑛)}, where 𝑛 is the total number of segments, consider each pair (𝑆𝑖, 𝑆𝑖+1). If the angle between segments, 𝜃𝑖, falls below a critical value, 𝜑, then the path is partitioned at the shared 𝜋 point, 𝑃𝑖. 𝑆𝑖 remains in the current partition, while 𝑆𝑖+1 starts a new one. Typically, a range of < 𝜑 < 2𝜋 is ideal for fitting curves, but this parameter must be tuned by the user.The relationship between 𝑆𝑖 and 𝑆𝑖+1 and the angle 𝜃𝑖 is graphically shown in Figure 5, andcan be expresses mathematically as:𝜃𝑖 = cos−1 𝑆⃑𝑖 ⋅ 𝑆⃑𝑖+1 |𝑆⃑𝑖| |𝑆⃑𝑖+1|(Eq. 1)where, 𝑆⃑𝑖 = 𝑃𝑖−1 − 𝑃𝑖 and 𝑆⃑𝑖+1 = 𝑃𝑖+1 − 𝑃𝑖.Figure 5: Calculating 𝜃𝑖 from segments6852Curvature PartitioningInitial partitioning results in several sets of line segments that need to be fit with parametric curves.

The curvature partitioning phase is the critical step of determining which line segments are best fit with arcs or splines. If a given partition contains a single line segment, no curve will be fit. The nature of parametric curves enables subsections to be classified based on the change in the discrete signed curvature of pairs of segments. Intuitively, discrete signed curvature describes how quickly consecutive line segments are changing their angle.Discrete signed curvature can be expressed mathematically as:1 𝑅𝑖𝑘𝑖 = 𝑠𝑖𝑔𝑛(𝜏𝑖) where, 𝜏𝑖 = 𝑆⃑⃑⃑(Eq. 2)𝑖 × 𝑆⃑⃑⃑𝑖+1 as a scalar, 𝑅𝑖 = |𝑃𝐶𝑖 − 𝑃𝑖|, and 𝑆𝑖, 𝑆𝑖+1 are not parallel.The definition of discrete signed curvature is rooted in finding the inverse radius Ri of a circle that passes through a set of three points. The signed component can be used to detect whether curvature is clockwise or counterclockwise.

The signed component can be used to detect whether curvature is clockwise or counterclockwise. It can be surmised that when curvature remains constant across a set of consecutive line segments, they are best fit with an arc. Conversely, when the curvature is dynamic, it is best fit with b-splines. Using this principle, subsections that are best fit with different curvature types are identified:Given a set of 𝑛 ordered segments {𝑆𝑖, (𝑖 = 0, … , 𝑛)} where 𝑛 > 1, each pair can be classified based on the change in their curvature ∆𝑘 = 𝑘𝑖+1 − 𝑘𝑖.𝑡𝑦𝑝𝑒 = {𝑎𝑟𝑐, ∆𝑘 ≅ 0 𝑠𝑝𝑙𝑖𝑛𝑒, 𝑜𝑡ℎ𝑒𝑟𝑤𝑖𝑠𝑒(Eq. 3)Graphically, the relationship between curve type and discrete signed curvature can be seen clearly. The examples below are cross-sections of triangulated meshes taking during the slicing process. The 3D models were exported from an implicit representation in modeling software to a triangulated mesh, discarding exact curvature information.

The 3D models were exported from an implicit representation in modeling software to a triangulated mesh, discarding exact curvature information. When the cross-section is taken, a series of line segments is generated approximating the surface of the mesh. Figure 6(a) shows the cross-section of a 3D model that was designed with three arcs and three line segments. Line segments were identified using the angle analysis in the initial partition phase. Note in Figure 6(b) that sections of line segments whose curvature is constant can be fit with arcs. The grouped segment points expressed in Figure 6(c) demonstrate how the series was partitioned.686(a)(b)(c)Figure 6: (a) Input model containing three arcs and three line segments, (b) discrete signed curvature of each consecutive pair of segments, (c) line segment points partitioned by curvature687Similarly, Figure 7(a) depicts a model that is constructed from two arcs and a b-spline.

When plotted in Figure 7(b), the two arcs are identified by grouping consecutive sections based on constant curvature. Since the other sections are dynamic, they are identified as a b-spline. Figure 7(c) demonstrates the results of the hybrid fitting algorithm as classified segment points.(a)(b)(c) Figure 7: (a) Input model containing a b-spline and two arcs, (b) discrete signed curvature of each consecutive pair of segments, (c) Line segment points partitioned by curvature688Even when preprocessing was applied, noise is still present in both above examples. This is the result of the Standard Triangle Language (STL) triangulation algorithm that is applied when models are exported from an implicit representation in CAD software. To compensate for noise in the discrete signed curvature data caused by triangulation, a threshold 𝛽 is used to determine if there was a substantial change in curvature.

To compensate for noise in the discrete signed curvature data caused by triangulation, a threshold 𝛽 is used to determine if there was a substantial change in curvature. Intuitively, 𝛽 is the maximum deviation from the mean of the change in curvature plus or minus 𝛿 times the standard deviation of the change in curvature. This statistical cutoff provides a method for filtering the noise generated by triangulation, while maintaining the classification capabilities when performing discrete signed curvature analysis. The value 𝛿 = 2 was used for this work; however, this parameter may need to be tuned for noisier datasets.Mathematically, 𝛽 is expressed as:𝛽 = 𝜇∆𝑘 ± 𝛿𝜎∆𝑘 (Eq. 4)where, 𝜇∆𝑘 =1𝑛𝑛 ∑ ∆𝑘𝑖 𝑖=1and 𝜎∆𝑘 = √∑(∆𝑘𝑖 − 𝜇∆𝑘)𝑛ResultsIn this section, we present several examples whose parametric curves were reconstructed from line segments using the proposed algorithm. The algorithm was implemented as a post- processing step into the ORNL Slicer 2.0 program [1].

The algorithm was implemented as a post- processing step into the ORNL Slicer 2.0 program [1]. This allowed for objects to be constructed with skins, shells, and a concentric infill pattern. Although the algorithm can be used with any set of polylines, concentric infill is the only pattern that will result in arcs or spline fits in Slicer 2. If selected, other infill patterns such as triangles or hexagons will not be fit with arcs or splines because their internal angles are below the threshold 𝜑. ORNL Slicer 2.0 additionally provided cross-sectioning and g-code exporting capabilities. Each model was designed with curvature in CAD software, exported as a triangulated mesh, cross-sectioned into layers, and then fit with arcs and splines. An analysis is provided into the toolpath file size reduction. Values of 𝜑 = , 𝛿 = 2, and 𝛼 = 10 were used.Example 1: A CylinderThe model in Figure 8(a) is a cylinder with a radius of 25 mm and a height of 100 mm.

Values of 𝜑 = , 𝛿 = 2, and 𝛼 = 10 were used.Example 1: A CylinderThe model in Figure 8(a) is a cylinder with a radius of 25 mm and a height of 100 mm. Given the nature of a cylinder, the model contains one circle per cross-section. From the toolpaths shown in Figure 9(a), it is observed that the cross-sectional outline is offset numerous times to form insets and skins. This results in multiple circles that need fit with G2 and G3 commands per layer. The arc-fit pathing in Figure 9(b) shows how each circle can be drawn with a single arc command. As indicated by Figure 8(a), the application of hybrid and arc-only fitting resulted in a reduction of motion commands by a factor of over 26 when compared with using exclusively line segments. Similarly, the application of spline-only fitting resulted in a reduction by a factor of 13. This epitomizes the situation where arc-only fitting is superior to spline fitting.

Similarly, the application of spline-only fitting resulted in a reduction by a factor of 13. This epitomizes the situation where arc-only fitting is superior to spline fitting. In this case, hybrid fitting produces the same result as arc-only fitting.6893𝜋4(a) (b)Figure 8: (a) input cylinder model, (b) number of motion commands required to construct the object(a) (b) Figure 9: (a) toolpaths using line segments only, (b) toolpaths after hybrid fitting, yellow sections show a single segment.Example 2: NURBS with HoleSeen in Figure 10(a), this example is a model designed with a continuous non-uniform rational basis spline (NURBS) around the outside, and a circular hole removed from inside. In Figure 11(a), the cross-sectioned NURBS curve results in a series of line segments best fit with b- splines. The circular hole will result in line segments that are best fit with a single arc command.

The circular hole will result in line segments that are best fit with a single arc command. Using the proposed method, the b-spline and arc segments were identified correctly and fit using their corresponding g-code commands. A major benefit of hybrid fitting can be realized in figure 10(b) by comparing the total number of motion commands required to construct an object. Arc fitting on this model provides a modest reduction in motion commands by a factor of 1.9 when compared with only line segments. Spline fitting improves this to a factor of 4.7 and hybrid fitting performs the best with a reduction of 6.8.

Spline fitting improves this to a factor of 4.7 and hybrid fitting performs the best with a reduction of 6.8. This model highlights the benefit of hybrid fitting over single-curve-type based approaches.690(a) (b)Figure 10: (a) input model drawn from a NURBS surface and circular hole, (b) number of commands required to construct the object(a) (b)Figure 11: (a) toolpaths using line segments only, (b) toolpaths after hybrid fitting691Example 3: 3D BenchyFigure 12 is the popular AM benchmarking model: 3DBenchy.stl. The boat is comprised of multiple sections that can be hybrid fit to reduce motion commands. Figure 13 shows a layer of toolpaths before and after hybrid fitting. Notable sections include: the bow section was replaced with splines, the small circle on the stern was replaced with circle arc commands, and several rounded corners were replaced with arcs.

Notable sections include: the bow section was replaced with splines, the small circle on the stern was replaced with circle arc commands, and several rounded corners were replaced with arcs. Using hybrid fitting cut the number of motion commands needed to construct the boat by 50% compared to only using line segments. Arc-only and spline- only fitting resulting in a reduction in commands by 27% and 41% respectively.(a) (b)Figure 12: (a) input 3DBenchy.stl, (b) number of commands required to construct the object(a) (b)Figure 13: (a) object built with only line segments, (b) object build with hybrid fit692Impact on Slicing TimeIn addition to recording the total number of toolpaths for each case study, wall-clock time was also measured for each case study. Table 1 demonstrates that the proposed hybrid curve fitting algorithm does not majorly impact the time required to slice and visualize an object.

Table 1 demonstrates that the proposed hybrid curve fitting algorithm does not majorly impact the time required to slice and visualize an object. In the Cylinder case study, the tool-path generation phase was slower due the added overhead of running the algorithm. However, the reduction of motions commands after this phase compared with no fitting results in faster optimize, write G-Code, and visualization steps. This contributes to hybrid fitting having a faster total slicing time when compared with no fitting. In the NURBS case study, a similar relationship is observed. In this case the total times are nearly identical. The hybrid fitting on the 3DBenchy ends up marginally slowing the slicing process. These examples suggest a correlation between the number of line segments replaced with curvature and a speed-up in processing time.

These examples suggest a correlation between the number of line segments replaced with curvature and a speed-up in processing time. This leads to the conclusion that even in cases where there is a modest reduction in motion commands, the algorithm presents only a marginal impact on slice time.ModelCylinderNURBS3DBenchyFittingNone Hybrid None Hybrid None HybridTool-path Generate (s) 1.3 7.1 0.9 3.0 3.1 6.2Optimize (s)0.6 0.2 0.1 0.1 0.3 0.3Write G-Code (s) 5.2 1.0 1.6 0.3 2.4 1.3Visualize (s) Total (s)3.0 0.5 0.9 0.2 1.6 0.910.0 9.0 3.5 3.6 7.4 8.6Table 1: Wall-clock time of each phase in the slicing workflow for case studies using no fitting and hybrid fittingConclusion and Further WorkA hybrid method for arc and spline reconstruction is presented in this paper for AM toolpath planning. It can dramatically reduce the number of required motion commands to construct an object with curved surfaces.

It can dramatically reduce the number of required motion commands to construct an object with curved surfaces. This reduction results in smaller g-code files and a better surface appearance of additively manufactured parts. Combining dense sections of line segments reduces choppy motion, thereby enabling machines with limited computational kinematic throughput to print more complex objects. Furthermore, we demonstrated that the proposed hybrid fitting can provide a tangible reduction in the number of motion commands over arc- and spline- only approaches to the parametric curve fitting problem. Leveraging discrete signed curvature analysis for hybrid fitting provides a robust method for classifying line segments in paths with dynamic geometry.

Leveraging discrete signed curvature analysis for hybrid fitting provides a robust method for classifying line segments in paths with dynamic geometry. Applying this method as a post-processing step enables seamless integration into existing slicing workflows, avoiding the complexities of developing a purely implicit toolpath planning program.Although the hybrid method is tolerant towards models with minimal noise, further investigation into fitting triangulated meshes that are reconstructed from noisy scan data is needed. W. Li et al. proposed using a low-pass filter to preprocess noisy scan data with their adaptive b- spline knot placement algorithm [8]. Perhaps, the same principle could be applied to noisy, reconstructed meshes. The outlined method also performs better when the input mesh is693triangulated with a uniformly high refinement. More investigation is needed into data preprocessing to provide better accuracy with paths that are sparsely populated.

More investigation is needed into data preprocessing to provide better accuracy with paths that are sparsely populated. Although the proposed method enables the implicit representation of printed geometry, it still relies on approximating CAD models with triangulated meshes as an intermediate step. This introduces the potential for slight deviations in printed object geometry when compared with the purely implicit CAD representation. The benefits of using the algorithm with curve-based infill such as the gyroid pattern could also be investigated. Finally, as more implicit geometry AM tools are developed, this tool could provide an interesting method for converting existing toolpaths into solid implicitly represented parts.AcknowledgmentsThis material is based upon work supported by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Office of Advanced Manufacturing, under contract number DE-AC05-00OR22725.References1.

BRID DEPOSITION AND MICRO ROLLING MANUFACTURING METHOD OF METALLIC PARTSHaiou Zhang*a, Yang Xiea, Daoman Ruia, Guilan WangbaState Key Laboratory of Digital Manufacturing Equipment and Technology, HuazhongUniversity of Science and Technology, Wuhan 430074, PR ChinabState Key Laboratory of Materials Processing and Die & Mould Technology, HuazhongUniversity of Science and Technology, Wuhan 430074, PR ChinaAbstractTo conquer the bottleneck problems of cracking, deformation, low accuracy and performance in existing additive manufacturing (AM) process, a new hybrid direct manufacturing method of metallic parts which integrates freeform deposition with micro rolling is proposed in this paper. The principle, devices and simulations are introduced and the comparison with Freeform Deposition Fabrication method are made that hybrid manufactured parts has distinctive features of higher accuracy and better microstructures.

The tensile strength is increased by 33% for stainless steel while the elongation percentage is improved more than 2 times. By using the Φ1.6mm wire in a feed rate of 1060mm/min, the deposition rate can get to 10kg/h. This method is appropriate for fabricating large-scale metal parts with outstanding quality, efficiency and low cost. Keywords: New AM method, HDMR (hybrid deposition and micro rolling), metallic parts, High performance, efficiency and low costthrough experiments, showing1. IntroductionWith the increasing of manufacture competition, developing an advanced manufacturing technology, which is efficient and low-cost, is particularly important. Freeform Deposition Fabrication (FDF) as a modern advanced manufacturing technology came into being.

Freeform Deposition Fabrication (FDF) as a modern advanced manufacturing technology came into being. However, as FDF is used in metal deposition, some technical bottleneck problems must be took into account, such as cracking, deformation, low accuracy and performance.The deposition layer surfaces of FDF parts are mostly out of flatness and there is always an error between the actual deposition thickness and the theoretical deposition thickness, which increases along the deposition height. The accuracy is not as high as that of traditional manufacturing. During FDF process, metal parts have to suffer a periodic procedure of rapid heating and cooling, which probably cause components deformation and cracking. Zhang et al. studied the freeform deposition _____________________Corresponding author E-mail: zholab@mail.hust.edu.cn (Haiou. Zhang). wgllab@mail.hust.edu.cn.conditions such as manufacturing path, energy power, traverse speed and cooling methods (Zhang et al., 2003).

Zhang). wgllab@mail.hust.edu.cn.conditions such as manufacturing path, energy power, traverse speed and cooling methods (Zhang et al., 2003). Under appropriate deposition conditions, components formability can be improved, but the problems of deformation and residual stress are still not fully resolved during thin-wall parts deposition process.To improve the accuracy of FDF method, some scholars measure the top surface morphology of the deposition layer to modify the deposition of the next layer. But it will reduce the manufacturing efficiency. Others combine traditional manufacturing method with FDF. Akula and Karunakaran, 2006 and Xiong, X.H., Zhang, H.O. et al., 2009 mill the part after deposition. Zhang.et al., 2010 proposes a hybrid deposition and micro rolling manufacturing method to significantly improve manufacturing accuracy, efficiency, formability and properties.

Zhang.et al., 2010 proposes a hybrid deposition and micro rolling manufacturing method to significantly improve manufacturing accuracy, efficiency, formability and properties. This method realizes the freeform deposition and the micro continuous rolling integrated in one manufacturing unit and makes the two processes work simultaneously. Components can be fabricated with high quality and high efficiency. So this new method is a solution to the bottleneck problems in the direct manufacturing field.Similar to the hybrid process, Contour Crafting (CC) is another hybrid rapid additive fabrication process (Khoshnevis et al., 2001). In order to create smooth and accurate planar and free-form surfaces of the object, it adopts the bladed towel instead of the roller to smooth out the striations between consecutive layers under computer control. Therefore, hybrid process is more suitable for metallic materials while CC is more disposed for non-metallic materials such as plastic materials.2.

Therefore, hybrid process is more suitable for metallic materials while CC is more disposed for non-metallic materials such as plastic materials.2. Principle, devices and simulations2.1 The principle and devices of hybrid manufacturing methodThe principle is shown in Fig.1. In hybrid manufacturing method process, metal parts are fabricated by hot rolling (the temperature is above recrystallization temperature) immediately after deposition. Comparing to FDF process, this method have three advantages. Firstly, the top surface flatness gets significant improvement, from middle-convex to nearly plane. What’s more, the residual stress can be reduced under a certain pressure. And the last one is that hybrid method can decompose the coarse grain into fine grains, eliminating cracking and deformation.Fig. 1. The principle of hybrid manufacturing method2.1.1 Energy source followed micro continuous rolling deviceAs shown in Fig.

1. The principle of hybrid manufacturing method2.1.1 Energy source followed micro continuous rolling deviceAs shown in Fig. 2, the energy source followed micro continuous rolling device integrates the energy source and the micro roller together. The roller would follow the energy source and smooth the top surface in deposition process. Even when the deposition path is curve or other complex path, the roller can fix on the top surface by rotating the part. Since the manufacturing temperature is very high, a water cooling system is designed to protect the entire device.Paul et al., (2013) also uses hybrid high-pressure rolling to deal with the deposition parts. But the improvements of accuracy and performance are limited since the rolling position of the device is far away from deposition position. Parts are not fabricated by hot rolling immediately after deposition. What’s more, their device can only be used in linear deposition.

Parts are not fabricated by hot rolling immediately after deposition. What’s more, their device can only be used in linear deposition. Through the comparison we can find our hybrid method is more flexible and space-saving since it can be used for manufacturing curved surface in one workstation.Cooling water pipeMicro rollerWelding gunFig. 2. Energy source followed micro continuous rolling device2.1.2 Deposition width adjustable micro continuous rolling deviceConsidering that some large-scale parts have broad-area slices, a width adjustable device is added to improve the deposition efficiency.

In this way the device can produce a larger deposition width, in the range of 0 to 30 mm, by means of the energy source swinging along the direction perpendicular to the forward direction.2.2 Simulations of hybrid manufacturing devicesSince the stress and strain field in the deposition manufacturing process is difficult to be measured, it is necessary to make process simulations for the hybrid manufacturing devices. And the results reveal the important role of hybrid manufacturing devices.2.2.1 Stress field simulationsThe maximum principal stress field simulations are shown in Fig.3. The deposition material is stainless steel and the roller is made of 45# steel. The initial temperature of the part is 293K while that of the roller is 313K. The deposition current is 220A and the voltage is 28.6V with 70% thermal efficiency. The part is 90mm long, 8.5 wide and 10mm height. The roller diameter is 30mm and the length is 30mm. Moving speed of the energy source is 0.01m/s.

The part is 90mm long, 8.5 wide and 10mm height. The roller diameter is 30mm and the length is 30mm. Moving speed of the energy source is 0.01m/s. The rolling reduction is 0.5 mm and the energy-roller distance is 50mm.the maximum principal stress field of FDF parts. The maximum principal stress moves from bottom of the parts to the top surface, which is of great effect for reducing the warping deformation of the base plate and leads to the improvement of manufacturing accuracy in deposition process.Thehybrid manufacturing method makesanupside-downchange(a) (b)Fig. 3. Maximum principal stress（Pa） field simulations of parts made by FDF (a) and hybrid manufacturing method (b)toWhat’s more, the maximum principal stress fields of the faces where the roller last leaves (the faces near us) are analyzed. The stress distributions are shown in Fig.4. After hybrid manufacturing, the original tensile stress of FDF parts decreases or even changes into compressive stress.

The stress distributions are shown in Fig.4. After hybrid manufacturing, the original tensile stress of FDF parts decreases or even changes into compressive stress. Compressive stress has a tendency to compensate for deformation, which is of great help to the inhibition of cracking and deformation.(a) (b)Fig.4. Width direction (a) and depth direction (b) maximum principal stress of the faces where the roller last leaves2.2.2 Strain field simulationsThe strain field simulations of FDF part and hybrid manufacturing part are shown in Fig.5-6. As is known to all, metal parts would contract with cold after deposition, which is a key problem to affect the dimensional accuracy. With hybrid manufacturing method, the strain values in both sides of metal parts increases, meaning the tensile deformations happen. These deformations compensate the cooling deformation, reducing the impact of cooling on the metal parts to some extent.(a)(b)Fig. 5.

These deformations compensate the cooling deformation, reducing the impact of cooling on the metal parts to some extent.(a)(b)Fig. 5. Maximum principal strain field simulations of parts made by FDF (a) and hybrid manufacturing method (b)(a) (b)Fig. 6. Width direction (a) and depth direction (b) maximum principal strain of the faces where the roller last leaves2.2.3 Temperature field simulationTo research on the microstructure of the deposition parts, temperature field simulations are necessary. As shown in Fig.7, because of the cooling system of the roller (the location of the black vertical lines) in hybrid manufacturing process, the heat-affected zone of the part made by hybrid manufacturing method is much smaller than that of the part made by FDF method. It means the material would be cooled faster and the grains would be smaller since they would stop grows after a certain temperature. The temperature curves of the parts simulated in Fig.7 are shown in Fig.8.

The temperature curves of the parts simulated in Fig.7 are shown in Fig.8. The temperature of the part made by hybrid manufacturing method plummets during the cooling of the roller, which contributes to the smaller grain size and better performance.(a)(b)Fig.7. Temperature simulations of parts made by FDF (a) and hybrid manufacturing method (b)Fig.8. Surface temperature curves of parts made by FDF and hybrid manufacturing3. Experiments of the hybrid direct manufacturing process3.1 ConditionsA 3-axis Computer Numerical Control (CNC) machine is employed to fabricate a linear thin-wall part. The stroke of X is 230 mm and that of Z is 160 mm. The raw material is stainless steel and plain carbon steel wire with 1.2 mm diameter. A Metal Inert-Gas welding gun is set and a micro continuous rolling device following the welding gun is used to realize hybrid manufacturing process.

A Metal Inert-Gas welding gun is set and a micro continuous rolling device following the welding gun is used to realize hybrid manufacturing process. The special G codes, consisting of the starting and ending instruction of deposition, are generated from self-developed software to make the hybrid process controlled automatically.3.2 ProcessThe manufacturing experimental process and temperature fields during the process are shown in Fig. 9. The high-temperature zone of part in FDF process is much larger than that of part in hybrid manufacturing process. And the temperature of the part after cooling of the roller is much lower than that in FDF process. In order to verify the accuracy of the temperature simulations above, the temperature of part made by hybrid manufactuing method is shown in Fig.10.(a) (b)Fig.9. Manufacturing experimental process of FDF method (a) and hybrid manufacturing method (b)(a)(b)Fig. 10.

Manufacturing experimental process of FDF method (a) and hybrid manufacturing method (b)(a)(b)Fig. 10. Temperature fields of manufacturing experimental processof FDF method (a) and hybrid manufacturing method (b)3.3 Experimental results3.3.1 AccuracyAs shown in Fig.11, the middle-convex top surface which occurs in FDF process would be flat after hybrid deposition. The problem of surface undulation and cracking has been solved effectively. And the problem of porosity in most deposition method can be suppressed to some degree. What’s more, the flowing phenomenon is alleviated and material utilization rat becomes higher. Above all, the thickness of each layer constant (1.2mm in the experiment), which means the thickness can be controlled accurately. That could be a breakthrough to solve the low accuracy problem of FDF brought by the error between the actual deposition thickness and the theoretical deposition thickness.3.3.2 Structure and propertiesAs shown in Fig.

12, the grain size of hybrid manufactured part is smaller. The density is higher and the microstructure is more homogeneous. A contrast tensile test (Fig. 13) was carried out to evaluate mechanical properties. For plain carbon steel, the tensile strength is increased by 33% and the tensile deformation is improved more than 2 times for hybrid manufactured part than those for FDF part.The same as the simulations, we can conclude from the result that the hybrid manufacturing method is capable to refine the microstructure and improve the tensile properties significantly. There might be two reasons why those happen. One clear reason is the utilization of water cooling system since the grain growth will be subjected to restraining.

There might be two reasons why those happen. One clear reason is the utilization of water cooling system since the grain growth will be subjected to restraining. Micro hot rolling may be another considerable factor since it can decompose the coarse grain into fine grains and transform as-cast microstructure into deformed microstructure.It is worth noting that there are two cooling sources for each layer during hybrid manufacturing process. One is the roller and another is the substrate for the first layer and previously deposited layers for the other layers. In general, the first layer has highest cooling rate because cold roller and cold substrate can absorb heat quickly. As with any other layered manufacturing process, there is a significant heat accumulation since layers (except for the first layer) are deposited on the previous layer which is already hot. In addition, surfaces of the roller have also been hot in subsequent rolling.

In addition, surfaces of the roller have also been hot in subsequent rolling. Thus, the cooling rate decreases along the deposition height. In order to maintain almost equivalent cooling rate for each layer, the suitable roller position should change as the build progresses.(a) (b)(c) (d)Fig. 11. Surface and profile comparison between parts made by FDF (a, c) andhybrid manufacturing method (b, d)Fig. 12. Optical microscope microstructure comparison between parts made by FDF (a) andhybrid manufacturing method (b)Fig. 13. Tensile test results comparison between parts made by FDFand hybrid manufacturing method4. The manufacture of a large-scale metallic part using hybrid manufacturing method4.1 Manufacturing conditionsThe thin-wall part with a complicated surface locally similar to the saddle could be fabricated by direct deposition manufacturing method. The 3-D CAD model of the part represented in STL format is shown in Fig. 14.

The 3-D CAD model of the part represented in STL format is shown in Fig. 14. A 6-DOF industrial robot is used in the manufacturing process. The welding gun and the hybrid manufacturing device are mounted on the end-effector of the robot. The raw material is welding wire with 1.6 mm diameter. As the width of each slice of the metal part, namely wall thickness of the part, is beyond the maximum width of single-pass deposition, the to realize hybrid deposition width adjustable micro continuous rolling device manufacturing.is usedFig. 14. STL modle of the thin-wall part4.2 Manufacturing processThe STL model of the part is decomposed into a series of slices by intersecting itself with a set of horizontal planes under a certain tolerance parameter such as the maximum allowable cups height. At the same time, the hatching paths for each slice are generated by appropriate process parameters such as hatching width and the pattern of hatching. The above procedures occur in the computer.

At the same time, the hatching paths for each slice are generated by appropriate process parameters such as hatching width and the pattern of hatching. The above procedures occur in the computer. Then the hatching paths represented in JBI format are sent to the fabricator, namely Motoman robot. Similar to the G code above mentioned, a JBI extension that consists of the starting and ending instruction of the MIG welding machine is applied. The robot reads and identifies a JBI file, then make the end-effecter move along the defined paths. Since this part is large-scale, lots of heat would be generated and built up during the manufacturing process and the water cooling method was used to reduce thermal deformation of the substrate. To make sure there was no porosity in the part, the shield gas was enable before the manufacturing starting and disable after the manufacturing ending. The manufacturing parameters are demonstrated in Table 1.Table 1.

The manufacturing parameters are demonstrated in Table 1.Table 1. The hybrid manufacturing parameters of the large-scale metal partEnergy powerEnergy sourceEnergy source-rollerEnergy source swingingWire speed(kW)speed (mm/min)distance (mm)frequency (Hz)(m/min)10450200.835.04.3 Manufacturing resultsAs shown in Fig.15, the maximum height of the part is 1200 mm, the maximum width is 1600mm and the average wall thickness is 15 mm. There is no obvious flowing phenomenon. Air-tightness test shows there is no hot cracking and porosity inside the part. The manufacturing accuracy is improved significantly, and by using Φ1.6mm wire in a wire feed rate of 1060mm/min in the fabricating process, the deposition rate of hybrid manufacturing method can get to 10kg/h.Fig. 15. The finished part5. ConclusionsThis paper proposed the HDMR manufacturing method and developed hybrid manufacturing devices.

15. The finished part5. ConclusionsThis paper proposed the HDMR manufacturing method and developed hybrid manufacturing devices. Simulations and experiments were carried out for the hybrid manufacturing process. The results and conclusions are as follows:(1) The top surface of the hybrid manufacturing deposition layers is flat instead of middle-convex and the profile has no cracking and porosity. The tensile strength is increased by 33% for stainless steel while the elongation percentage is improved more than 2 times.

The tensile strength is increased by 33% for stainless steel while the elongation percentage is improved more than 2 times. So this new method can reduce flowing phenomenon and improve the components formability and accuracy.(2) The deposition layers’ thickness can be controlled accurately using this method, which can resolve the common problems in present FDF technology brought by the error between the actual deposition height and theoretical deposition height.(3) The simulations of stress, strain and temperature fields show that the hybrid manufacturingmethod makes contributions to the more compact microstructure with smaller grains.(4) A large-scale metal part is fabricated using this new method, which shows the hybrid process is capable to fabricate large-scale metal parts with low cost, high quality and great efficiency.

By using Φ1.6mm stainless steel wire in a wire feed rate of 1060mm/min in thefabricating process, the deposition rate of hybrid manufacturing method can get to 10kg/h, which is much more than LENS. So it has great value for industrial application.So far, this hybrid manufacturing method has been utilized with top surface rolling only,expanding it into top and side surface rolling at the same time is the trend for this new method.AcknowledgmentsThe authors would like to thank Juqiu Gong, Feiyong Wang, Yang Zhang, Pengyang Xiang, Yanzhao Peng and Qi He for their assistance with this study. In addition, the contribution of the National Nature Science Foundation of China under Project No. 51175203 is gratefully acknowledged.ReferencesPaul, A., Colegrove, Harry E., Coules, Julian, Fairman, Microstructure and residual stress improvement in wire and arc additively manufactured part through high-pressure rolling.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperHybrid Geometry/Property Autoencoders for Multi-Lattice Transitions Martha Baldwin,1 Nicholas A. Meisel,2 Christopher McComb11Department of Mechanical Engineering, Carnegie Mellon University, Pittsburgh, PA 15213 2School of Engineering Design and Innovation, The Pennsylvania State University, University Park, PA 16802Abstract Additive manufacturing has revolutionized structural optimization by enhancing component strength and reducing material requirements. One approach used to achieve these improvements is the application of multi-lattice structures. The performance of these structures heavily relies on the detailed design of mesostructural elements.

One approach used to achieve these improvements is the application of multi-lattice structures. The performance of these structures heavily relies on the detailed design of mesostructural elements. Many current approaches use data-driven design to generate multi-lattice transition regions, making use of models that jointly address the geometry and properties of the mesostructures. However, it remains unclear whether the integration of mechanical properties into the data set for generating multi-lattice interpolations is beneficial beyond geometry alone. To address this issue, this work implements and evaluates a hybrid geometry/property machine learning model for generating multi-lattice transition regions. We compare the results of this hybrid model to results obtained using a geometry-only model.

We compare the results of this hybrid model to results obtained using a geometry-only model. Our research determined that incorporating physical properties decreased the number of variables to address in the latent space, and therefore improves the ability of generative models for developing transition regions of multi-lattice structures.1. IntroductionAdditive manufacturing has enabled more design freedom, however, these freedoms have forced designers to become more creative due to continuous pushes for optimization and material reduction. One approach to keep pace with these pushes is to utilize lattice structures, which can be used to reduce weight while maintaining performance [1,2]. By simply applying lattice structures to a design, weights have been shown to decrease by as much as 40% while maintaining overall strength [3].

By simply applying lattice structures to a design, weights have been shown to decrease by as much as 40% while maintaining overall strength [3]. This is one of many benefits that have sparked further interest in lattices, leading to the development of multi-lattice structures which are structures created using multiple types of unit cell topologies [4–7]. Multi-lattice design has become a major area of interest among researchers in additive manufacturing due to its versatility in terms of properties. Of note, multi- lattice structures exhibit better strength and stiffness properties than uniform structures of comparable density [6,8].However, the success of multi-lattice structures is dependent on their ability to distribute stress evenly, as stress concentrations can cause part failures [6]. Due to the complex nature of this design problem, researchers have explored approaches that use machine learning to design multi-lattice transition regions [9–13].

Due to the complex nature of this design problem, researchers have explored approaches that use machine learning to design multi-lattice transition regions [9–13]. These models often make use of both geometry and stiffness information in the training data to create generative models that are responsive to both. Stiffness is a physical property that is considered in most of the literature that analyzes the performance of lattices [2,14,15], as it is a primary physical descriptor of the performance of a lattice. While incorporating stiffness properties into these models has proven successful, the relative value of incorporating1941physical properties into these generative models has not been evaluated [9–13].

In other words, it is unclear whether the added complexity and computational requirements incurred through the addition of stiffness information provide an associated increase in design performance to justify those costs.This study aims to determine whether it is necessary to incorporate physical properties into variational autoencoders (VAEs) that use geometrically-defined latent spaces to optimize the physical performance of transition regions of multi-lattice structures. Specifically, we examine stiffness continuity, a measure of the change in stiffness across adjacent unit cells, as an indicator of the stiffness throughout a transition region. Our primary research questions address the unknowns regarding the role of physical properties in unit lattice cell latent spaces:1. How does incorporating physical properties into VAEs affect the relationship between geometric smoothness, distance, and transition length in the latent space?2.

How does incorporating physical properties into VAEs affect the relationship between geometric smoothness, distance, and transition length in the latent space?2. How does incorporating physical properties into VAEs affect the relationship between stiffness continuity, distance, and transition length in the latent space?2. MethodologyThis section outlines the architectures developed and the methods of testing used to evaluate our VAEs. Additionally, it will briefly outline our use and development of data for this work.The data used was originally generated in prior work by Wang et al. and consists of 248,396 unique orthotropic microstructures, or unit cells [10,12,16]. In addition to this data, we developed a function for calculating the stiffness tensors of generated geometries based on a MATLAB topology optimization code [17]. This code was validated against those computed by Wang et al in prior work [10,12,16].

This code was validated against those computed by Wang et al in prior work [10,12,16]. A subset of 10,000 random data points was used to develop architecture of the models without excessive training times. The individual pieces of data were binary arrays of size 50×50 that represented an individual unit cell’s geometry. Additionally, each unit cell had a corresponding 3×3 stiffness tensor.2.1 Architectures This section outlines the VAE architectures used for investigating the research questions. Specifically, this entails two main architectures: (1) a geometry-only VAE based on prior work, and (2) a hybrid representation VAE that encodes both stiffness and geometry information. Both architectures in this work used identical training parameters: training using a batch size of 32 and an Adam optimizer [18], where 85% of the dataset was used for training.

Both architectures in this work used identical training parameters: training using a batch size of 32 and an Adam optimizer [18], where 85% of the dataset was used for training. Additionally, training was terminated early if the loss failed to improve after 10 full epochs, where the loss term measures the difference between the reconstructed data and the original data.The geometry VAE simply encodes and decodes the geometry using a standard VAE architecture (see Figure 1). The architecture of the model in Figure 1 was based on our previous works [19,20]. The purpose of this model is to serve as a baseline for comparison against the hybrid architecture.1942Figure 1: Geometry VAE Architecture* where the input consists of only the unit cell geometryThe hybrid VAE (see Figure 2) has a similar encoder and decoder framework to the geometry VAE. However the encoded geometry is appended with a flattened stiffness matrix before calculating the mean and log variance.

However the encoded geometry is appended with a flattened stiffness matrix before calculating the mean and log variance. This architecture is intended to encode information relating to the stiffness tensor into the latent space. The decoder of this architecture is identical to the decoder of the geometry decoder architecture to enable a consistent comparison.Figure 2: Hybrid VAE Architecture* where the input consists of unit cell geometry and the corresponding stiffness tensor2.2 Evaluating Performance of Architecture The geometry VAE and hybrid VAE must be evaluated in terms of both geometric smoothness and stiffness continuity. Prior work establishes a procedure for evaluating geometric smoothness and shows that it is related to the distance in latent space and transition length [19,20]. We use the same analysis procedure and geometry smoothness metric here.

We use the same analysis procedure and geometry smoothness metric here. Specifically, distance is measured with respect to number of standard deviations, and transition length is simply the number of points in a transition region. We also introduce a new metric for evaluating stiffness continuity.2.2.1 Evaluating Geometric Smoothness The success of a latent space is dependent on the performance of the interpolations that can be produced from that latent space. When evaluating geometric transitions in a latent space, continuously and smoothly changing geometries are desired. To perform this evaluation, we utilize a smoothness metric that was developed in our previous work to evaluate the geometric smoothness of a 2D interpolation [19,20]. This metric calculates the gradients between multiple layers in an interpolation, essentially measuring the flow between each layer of pixels.

This metric calculates the gradients between multiple layers in an interpolation, essentially measuring the flow between each layer of pixels. The root mean squared error (RMSE) is calculated between the flows and then normalized to produce aImage created with http://alexlenail.me/NN-SVG/AlexNet.html1943value of smoothness over an entire interpolation. This metric will be utilized in this work to evaluate the smoothness of interpolations among the various models. More details on the implementation of this smoothness evaluation are available in prior work by the authors [20].2.2.2 Evaluating Stiffness Continuity When evaluating the equivalent of geometric smoothness for stiffness tensors, the changes within a single tensor are not important as each element represents a unique piece of information. Therefore, the tensors are evaluated by comparing values only with the neighboring tensor. To achieve this, we implemented an RMSE framework to calculate a stiffness continuity value.

Therefore, the tensors are evaluated by comparing values only with the neighboring tensor. To achieve this, we implemented an RMSE framework to calculate a stiffness continuity value. First, the stiffness tensors in the interpolation were normalized with respect to the entirety of the training data. Then, the RMSE must be computed between each pair of values in the stiffness tensor using𝑅𝑀𝑆𝐸! = 	 ’& ∑ )𝐾!"#,% − 𝐾!,%, 𝑁’ %(#( 1 )where 𝑅𝑀𝑆𝐸! denotes the root mean squared error between a pair of stiffness tensors at indices i and i+1, K is the normalized flattened stiffness tensor, j is the index that identifies the specific term in flattened stiffness tensor, and N is the number of terms in a single stiffness tensor (in this case, 9). Since the stiffness tensors are normalized before computing the RMSE, the maximum possible value is 1.

Since the stiffness tensors are normalized before computing the RMSE, the maximum possible value is 1. This allows the direct evaluation of the continuity of the stiffness by averaging the values in Equation 1 using𝐶) = )1 − 𝑎𝑣𝑔(𝑅𝑀𝑆𝐸!), ⋅ 100%( 2 )where 𝐶) is a value representing the continuity of stiffness over the entire transition region. Higher values indicate a smoother and more continuous transition in stiffness, while lower values indicate abrupt transitions.3. ResultsIn this section we discuss results of the models in three main areas: (1) the reconstruction capabilities, (2) the test interpolations outlined in the methods, and (3) an ordinary least squares evaluation based on the interpolation results.3.1 Model Reconstruction Initially, we evaluate the performance of a machine learning model by examining the performance during training and testing (here using a mean squared error (MSE) loss and coefficient of determination).

This indicates whether or not the model can appropriately reconstruct the desired data. If the model can reconstruct data, then an examination of the learned latent space may be informative. The results from the geometry VAE (see Figure 3) showed that geometry can be effectively reconstructed to 80% accuracy for testing data, and 87% accuracy for training data. This model serves as the baseline and represents the original model developed in previous work [19,20]. The decrease in accuracy from our previous work is likely due to the increase in dimensionality of the data, as well as the randomness of the data. The performance of the validation1944data was likely extremely inconsistent due to the lack of similarity between all the data points.

The performance of the validation1944data was likely extremely inconsistent due to the lack of similarity between all the data points. This is a matter to consider when evaluating the other model types, as this model serves as the baseline.Figure 3: Geometry VAE: Plot of Loss and Coefficient of Determination (left) and Visualization of the Latent Space using PaCMAP Dimensionality Reduction (right)The hybrid VAE was nearly identical to the geometry VAE, with matching performance for accuracy for both training and testing data (see Figure 4).

This is a good indication that the model will perform well when reconstructing data, and the results in the evaluation section should be directly comparable.Figure 4: Hybrid VAE: Plot of Loss and Coefficient of Determination (left) and Visualization of the Latent Space using PaCMAP Dimensionality Reduction (right)In addition, Pairwise Controlled Manifold Approximation (PaCMAP) was used to visualize the latent space produced by each of the trained models (see Figure 3 and Figure 4) [21]. This approach has been shown in prior work to accurately provide a representation with balanced preservation of local and global features for engineering-relevant data [22]. There are distinct similarities in the learned embeddings, despite the addition of unique performance information. The overall shape is similar, consisting of approximately four lobes.

There are distinct similarities in the learned embeddings, despite the addition of unique performance information. The overall shape is similar, consisting of approximately four lobes. In addition, the unit cells with lower average density are placed at the periphery of the latent space.19453.2 Interpolation Performance As discussed previously, the evaluation technique for consistent results was traveling through latent space using standard deviations as the distance metric. This metric is discussed in further detail in prior work [19,20]. Each point on a plot represents an interpolation, with the average value of a metric (y-axis) over a distance in the latent space (x-axis). The points are labeled using color to denote the length of the transition region.The results from Figure 5 (left) are consistent with the results seen in previous work [19,20]. As distance in the latent space increases, the smoothness decreases.

As distance in the latent space increases, the smoothness decreases. Additionally, the higher transition lengths have higher smoothness values relative to one another.Geometry SmoothnessStiffness ContinuityFigure 5: Geometry VAE: Geometry Smoothness versus Number of Standard Deviations in the Latent Space (left) and Stiffness Continuity (𝐶)) vs Number of Standard Deviations in the Latent Space (right).The predicted geometries from Figure 5 (left) were used to the calculate the corresponding stiffness tensors, which were evaluated using the metrics described in the methods section to produce Figure 5 (right). This figure serves as the baseline of stiffness continuity for our hybrid geometry/property model. From visual observations, the stiffness continuity begins to plateau as the distance in the latent space increases.

From visual observations, the stiffness continuity begins to plateau as the distance in the latent space increases. This indicates that the effect of distance in the latent space is limited to some degree.The results from the hybrid VAE are consistent with the results from the geometry VAE, where the relationship between smoothness, distance, and transition length are comparable (see Figure 6 (left)). This is a desirable trait, as it indicates that the latent space has similar embeddings to the baseline model. Which means that we can directly compare the performance of the two models in terms of the stiffness continuity.

Which means that we can directly compare the performance of the two models in terms of the stiffness continuity. However, the pattern of the results displays a much more aggressive plateau starting at 4 standard deviations between endpoints.1946Geometry SmoothnessStiffness ContinuityFigure 6: Hybrid VAE: Geometry Smoothness versus Number of Standard Deviations in the Latent Space (left) and Stiffness Continuity (𝐶)) vs Number of Standard Deviations in the Latent Space (right).The predicted geometries from Figure 6 (left) were used the calculate the corresponding stiffness tensors, which were evaluated using the metrics described in the methods section to produce Figure 6 (right).

The stiffness continuity of the data in Figure 6 (right) appears to plateau much sooner than the points in Figure 5 (right).Visually, Figure 5 (left) and Figure 6 (left) are consistent with prior work, showing that smoothness has a clear relationship to distance and the length of the transition region. Smoothness appears to be positively correlated with the length of the transition region, and negatively correlated with the distance in the latent space between the endpoints. Figure 5 (right) and Figure 6 (right) show the relationship between stiffness continuity, distance in the latent space, and length of the transition region. These plots demonstrate that the effect of distance and transition length does not have as significant of an impact as they do on geometric smoothness. However, these values may be skewed arbitrarily high, given that symmetric geometries contain zeros in the stiffness tensor [23].

However, these values may be skewed arbitrarily high, given that symmetric geometries contain zeros in the stiffness tensor [23]. The next section serves to further explore the validity of these visual observations.3.3 Evaluating Latent Space Relationships This section will outline the relationships between the primary variables in the latent space. The independent variables are smoothness of geometry and continuity of stiffness properties of the respective geometries. The dependent variables are distance in the latent space and the length of the transition region. Using these variables, an ordinary least squares regression can be executed based on the results in the interpolation performance section.Table 1 displays the ordinary least squares regression analysis of the results from Figure 5 (left), which evaluates the relationships between geometric smoothness and latent properties for the geometry model.

Again, the geometry model serves as the baseline since it has architecture similar to that used in prior work [19,20].

This is evident based on the p-values in Table 1. The high R-squared value is indicative that the variability in the geometric smoothness is almost fully described by the two independent variables, distance and transition length.Based on the results of Table 1 and Table 2, the relationship between stiffness continuity and the properties of the geometry defined latent space are not as strong as the relationship with geometric smoothness and the properties of the latent space. This is based on the low R-squared value of the stiffness continuity of the geometry VAE in Table 2. An R-squared value of 76% indicates that 24% of the variability of the stiffness continuity is not accounted for by the two independent variables explored. This suggests that there are additional variables that affect the performance of stiffness continuity.

This suggests that there are additional variables that affect the performance of stiffness continuity. Therefore, the true underlying relationship between stiffness continuity, distance, and transition length is unclear.Table 3 displays the ordinary least squares regression analysis of the results from Figure 6 (left), which evaluates the relationships between geometric smoothness and latent properties for the hybrid model.

This conclusion is based on the p-values of the transition and interaction terms in Table 3. Alternatively, stiffness continuity in the hybrid model has a relationship with the transition length, based on the p-values from Table 4. However, the R-squared values of nearly 90% suggest that other variables account for approximately 10% of the variability of the geometric smoothness and stiffness continuity.Overall, the results indicate that incorporating physical properties into the VAE changed the relationships with distance and transition length. The results suggest that incorporating physical properties into the latent space decreased the likelihood that transition length influences geometric smoothness. However, stiffness continuity demonstrated a greater likelihood that the combination of distance and transition length have an effect. As expected, none of the models showed a relationship with transition length alone.

As expected, none of the models showed a relationship with transition length alone. The results from the hybrid VAE indicate that its latent space is better suited to create multi-lattice transitions with continuous geometry, given that fewer variables affect its performance. However, none of the results relating the two models were statistically significant, thus further testing is warranted.4. ConclusionThe development of multi-lattice transition regions is increasingly dominated by generative machine learning models. These models aim to produce multi-lattice transition regions with smooth geometric and physical properties. This work compares two approaches to constructing transition regions: a model that uses only geometric information and another hybrid model that merges geometry and property information. The first model, the geometry VAE, created a geometrically defined latent space only using the unit cell topologies.

The first model, the geometry VAE, created a geometrically defined latent space only using the unit cell topologies. The second model, the hybrid VAE, utilized a combination of unit cell topologies and unit cell stiffnesses to define the latent space.The geometry VAE performed much like our previous model, where geometric smoothness is influenced by distance and the combination of distance and transition length. Further testing is1949necessary to fully explain the relationships between stiffness continuity and the latent space. The hybrid VAE exhibited very different relationships than the geometry VAE. The results of the ordinary least squares regression indicate that incorporating stiffness into the generative model decreased the effect of transition length with respect to geometric smoothness. This is a desirable trait as it indicates that the geometric smoothness will only be affected by the distance in the latent space.

This is a desirable trait as it indicates that the geometric smoothness will only be affected by the distance in the latent space. Alternatively, the continuity of stiffness is affected by the combination of transition length and distance when analyzing the hybrid model. However, there appears to be a set of additional variables that account for ~10% of the variability for both stiffness continuity and geometric smoothness.When analyzing the reconstruction accuracy of the models, they can only perform at 80% accuracy, which is much lower than previous work. The drop in accuracy from our previous work is likely due to the increase in dimensionality of the data, from 28×28 to 50×50. Although the accuracy is sufficient for the current analysis, further work should seek to train more accurate VAE models. In addition, there are many ways in which property information can be combined with geometric information in the training process of hybrid models.

In addition, there are many ways in which property information can be combined with geometric information in the training process of hybrid models. Therefore, future work should also explore the design freedom that is present in the design space of deep learning models.AcknowledgementsThis material is based upon work supported by the National Science Foundation through Grant No. CMMI-1825535. Any opinions, findings, conclusions, or recommendations expressed in this paper are those of the authors and do not necessarily reflect the views of the sponsors. We are also grateful to Yigitcan Comlek from the Integrated DEsign Automation Laboratory (IDEAL) at Northwestern University for generously sharing their data to support this work.References[1]Hanks, B., Berthel, J., Frecker, M., and Simpson, T.

Preventing dielectric breakdown of the printed substrate, the ultrasonic embedding approach enables complex geometrical embedding through customized software. This work presents the effective embedding of large diameter (14 AWG) kapton coated litz wire into polycarbonate (PC) substrate. Custom software allowed for generation of embedding toolpaths directly from the CAD model of the designed coupon. Results showed the most successful embedding paths were circular pre-formed cavities. Through characterization of a myriad of printed samples, an approach was developed for embedding large diameter wire. Through the use of the Foundry Multi3D System, the increased complexity of embedded electronic parts can further impulse the implementation of hybrid additive manufacturing in large scale applications.1.

Introduction The definition of hybrid manufacturing has been well established as an amalgamation of two or more manufacturing processes to fabricate parts within a simplified process chain, and a condensed amount of time. In conventional manufacturing, work was performed by Jeng et al., (2001) in the development of a hybrid process by combining selective laser cladding (SLC) and three-axis milling. Hybrid processes are not exclusive to multiple machines, as the use of multiple cycles within one machine could achieve “hybridization”. Hur et al., (2002) developed a hybrid rapid prototyping process that implemented two cycles of operation. The first cycle involved front and back machining, while the second cycle implemented drilling, milling, and grinding for achieving the final part. Extending beyond conventional means, hybrid additive manufacturing (AM) allows for the fabrication of complex geometries due to collaborative processes.

Extending beyond conventional means, hybrid additive manufacturing (AM) allows for the fabrication of complex geometries due to collaborative processes. For example, Karunakaran et al., (2010) and Du et al., (2016) presented a hybrid system that made use of selective laser melting (SLM), an AM method, with the addition of milling, a subtractive process. Lee et al., (2014) developed a hybrid rapid prototyping system by integrating a Fused Deposition Modeling (FDM) nozzle, with a cutting spindle on a five-axis desktop machine. In polymer material extrusion AM, access to every layer during part fabrication has enabled an increase in component functionality. Multifunctionality refers to the ability of a component to be utilized beyond structural purposes. Espalin et al., (2014) described a method of adding electrical functionality by implementing wire embedding on FDM printed thermoplastic357substrates.

Espalin et al., (2014) described a method of adding electrical functionality by implementing wire embedding on FDM printed thermoplastic357substrates. Other work was performed to add electrical and magnetic functionality within plastics by introducing silver based conductive inks and solid copper wire [Shemelya et al., (2015)(2013); Joe Lopes et al., (2012)]. Research conducted at the W.M. Keck Center for 3D Innovation, in pursuit of hybrid manufacturing of multifunctional components, led to the the Foundry Multi3D development of System that incorporates ultrasonic wire embedding. Shown in Figure 1, the Foundry Multi3D System consists of two FDM Fortus 400mc (Stratasys, Eden Prairie, MN, USA) industrial printers, a computer numerical control (CNC) router (LC 3024 CNC), and a six-axis robotic arm (Yaskawa Motoman MH50, Miamisburg, OH, USA)(Ambriz et al. 2017).

2017). Through modification of the CNC router, additional processes were implemented to include the ability to machine, embed foils and wires, and pick and place electronic components (Coronel et al., 2017). Leveraging the ability to interrupt the printing process, a portable build platform was made to transfer the workpiece to the CNC for embedding, machining, etc., then be returned to the FDM printers. Overall, the Foundry Multi3D System facilitates the fabrication of 3D printed electronics.FDM1CNC Routerportable build platformsix-axis robotFDM2Figure 1: Foundry Multi3D SystemIntegrating electrical functionality to additively manufactured structures is possible through UTEP’s patent-pending wire embedding technology. Ultrasonic consolidation has been utilized in several applications, such as bond formation between aluminum matrix and SiC fiber Yang et al., (2009).

Ultrasonic consolidation has been utilized in several applications, such as bond formation between aluminum matrix and SiC fiber Yang et al., (2009). Although significant work has been performed in ultrasonic consolidation of metal matrix composites Sriraman et al., (2014), this method has been adopted for the ultrasonic welding of plastic as well Yousefpour et al., (2004). Hybrid AM was accomplished in this study, by adopting a commercially available ultrasonic welding tool on the CNC router of the Foundry Multi3D System. This allowed for the embedding of kapton coated litz wire, traditionally used for power electronics. Kapton coated wire permits the use of high currents, while preventing dielectric breakdown of the substrate. The work presented, describes the approach for embedding large diameter kapton coated wires for future use in power electronics component fabrication, including stators for electric motors.2.

The work presented, describes the approach for embedding large diameter kapton coated wires for future use in power electronics component fabrication, including stators for electric motors.2. Approach In general, ideal wire embedding characteristics include minimal part deformation, adhesion between the wire and substrate, and dimensional accuracy. Wire should be flush with the surface, as surface uniformity is essential for continued printing over the embedded substrate. Deformation would be evidenced by heat damage, displaced material, and physical warping of the part. Figure 2 shows a printed polycarbonate (PC) coupon upon which two sizes of kapton coated wires were embedded. The three wires on the right were 28 AWG wires, while the single wire on the left was a 14 AWG wire. The cross section of the coupon was observed with an OGP SmartScope Flash 250 (OGP, Rochester, NY). This initial experiment set the premise for the presented research.

The cross section of the coupon was observed with an OGP SmartScope Flash 250 (OGP, Rochester, NY). This initial experiment set the premise for the presented research. When the 28 AWG wire was embedded, PC material was found to be completely surrounding the 358displaced materialmaterial surrounding wireFigure 2: Embedded wires with cross-sectional views perimeter of the wire. Using previously described criteria, the 28 AWG wires were characterized as well embedded. For the 14 AWG wire, material displacement was observed, in addition to the wire veering from the intended embedding path. An approach was developed through the use of preformed cavities, to allow the proper embedding of large diameter kapton coated wire.

An approach was developed through the use of preformed cavities, to allow the proper embedding of large diameter kapton coated wire. Experimental approach and results are discussed, as well as the possible applications for this technology as it pertains to hybrid additive manufacturing and power electronics.For all experimentation, PC specimen were printed on the Foundry Multi3D System, where once fabricated, the robot arm would transfer them into the CNC for embedding. A commercially available ultrasonic welding tool was used for embedding. The ultrasonic vibrations of the tool heated the printed substrate, causing plastic deformation to allow for the wire to be introduced into the thermoplastic. Figure 3 shows a schematic of the ultrasonic wire embedding tool that was mounted on the CNC router. This tool utilized several horns, depending on the application.

Figure 3 shows a schematic of the ultrasonic wire embedding tool that was mounted on the CNC router. This tool utilized several horns, depending on the application. Although some horns were designed to simultaneously feed wire and secure it to the plastic, the horn used for these experiments was a flat tip horn that traversed over the manually laid wire. The wide range of embedded wire options required differing approaches to allow for proper embedding. Often, small diameter wires could be fed through a specific horn and embed into a flat substrate without significant struggle. This was attributed to the fact that material displaced by their introduction was minimal. For large diameter wire this was not the case. The material displaced by introducing large diameter wires significantly hindered the ability to evenly deposit material on the surface of the substrate. For this reason, cavities were introduced into the substrate design.

For this reason, cavities were introduced into the substrate design. The preformed cavities would allow for the wire to be guided, for improved dimensional accuracy, while embedded completely to allow for subsequent material deposition.ultrasonic horncopper wirePC substrateto understand the embedding process, and the effect of generated temperature between the tip and the substrate. Infrared thermographySeveraltests were performedCNC Router BedFigure 3: Ultrasonic embedding schematic359the was used embedding speed that would circumvent excessive melting and deformation of the plastic. A sample IR thermography result is shown in Figure 4. The figure captures the ultrasonic horn contacting the plastic substrate while ultrasonically embedding a 28 AWG wire. Localized heat generation was observed at the contact point, where the temperature reached an approximate range of 50-55C.

Localized heat generation was observed at the contact point, where the temperature reached an approximate range of 50-55C. The temperature was relatively the glass transition temperature of the polycarbonate substrate (~145C), let alone the melting temperature (~280C). These results proved that ultrasonic embedding does not rely specifically on heat generation, but rather plastic deformation of material, to allow the wire to be embedded.to aidin determiningultrasonic hornPC substratelow comparedtoFigure 4: IR thermography of ultrasonic embedding on PC3. Experiment Embedding low gauge wire into a plastic component proved to be a challenge. Material displacement due to large diameter wires, prevents subsequent material extrusion. Therefore, cavities were implemented to introduce wire in a feasible manner, without compromising the fabrication of the component.

Therefore, cavities were implemented to introduce wire in a feasible manner, without compromising the fabrication of the component. The cavity geometries were selected based on the simplest shapes that a circular wire could be introduced into. With the flexibility of 3D printing, four different cavity geometries were fabricated; a circle, a square, and two equilateral triangles (regular and inverted), shown in Figure 5. These cavities were designed using Autodesk’s Fusion 360 software and manufactured on a Fortus 400mc within the Foundry System. The coupons were all 1x1 inches square, and the 14 AWG kapton coated litz wire was used as the standard for all testing. Litz wire SquareTriangleInverted TriangleFigure 5.

The coupons were all 1x1 inches square, and the 14 AWG kapton coated litz wire was used as the standard for all testing. Litz wire SquareTriangleInverted TriangleFigure 5. Design and fabrication of preformed cavity360Table 1: Dimensional Accuracy for Printed SpecimensGeometrySample NumberDiameter (inch)GeometrySample NumberBase (inch)Height (inch)10.07310.0630.053Circle20.068Triangle20.0610.05230.07030.0590.048Length (inch)Base (inch)Height (inch)Square120.0580.057Inverted triangle120.0910.0940.0830.08730.05730.0920.089is composed of multiple wire strands, which made the diameter and area vary. The average diameter for the 14 AWG wire was 0.066 inches. Utilizing the cross-sectional area of the wire, the areas for the cavities were determined. The circular cavity was made 0.07 inches so that there would be enough space for the wire to be introduced.

Utilizing the cross-sectional area of the wire, the areas for the cavities were determined. The circular cavity was made 0.07 inches so that there would be enough space for the wire to be introduced. Based on the area of the circular cavity, the other cavities were designed. Another factor in cavity size, was the manufacturing capabilities of the printer. Layer thickness was set to 0.01 inches, which resulted in some cavities being slightly larger because the size was rounded up. Table 1 shows the dimensions of the three cavities that were printed per sample. Each sample was inspected with the OGP SmartScope, from which the dimensions were recorded.The experimentation involved using the flat ultrasonic horn on the Foundry System. Pauses were strategically placed such that the entirety of each geometrical cavity was formed. The specimen with the open cavities were transferred from the Fortus 400mc via robot arm, into the CNC router with the ultrasonic horn assembly.

The specimen with the open cavities were transferred from the Fortus 400mc via robot arm, into the CNC router with the ultrasonic horn assembly. The wires were manually laid over their respective cavity, and the horn traversed over each wire. Embedding the wire thermoplastic required more than just the ultrasonic tool. The CNC gantry system was prepared with a complimentary software add-in. To obtain accurate the embedding linesintothethroughoutFigure 6: Cavity embedding and results361Figure 7: Cross sectional resultsprocess, Fusion 360 was utilized to serve as a modeler. In Figure 6, the wire paths were added to the CAD and by using Fusion’s CAM package, the paths were processed. A custom post process plug-in was added to Fusion’s CAM to obtain a GCODE. This GCODE file was then used on the CNC router for wire embedding implementation.

A custom post process plug-in was added to Fusion’s CAM to obtain a GCODE. This GCODE file was then used on the CNC router for wire embedding implementation. The file includes parameters to turn on/off the ultrasonic tool, jog speed, and pauses for each circuitry.After testing, it was determined that a 60 mm/min jog speed for the traversing ultrasonic tool provided the best embedding. Figure 7 shows the cross section of each geometry at both the un-embedded and embedded state. As the figure shows, the circular cavity resulted in the wire with the least amount of deformation. The clear boundary line in the circular sample suggests a high level of wire-plastic contact throughout the circumference. This wire-plastic interface was not as pronounced in the other samples. The square cavity showed significant wire deformation, as well as gaps in the top two layers of the sample.

This wire-plastic interface was not as pronounced in the other samples. The square cavity showed significant wire deformation, as well as gaps in the top two layers of the sample. The inverted triangle sample had acceptable wire-plastic contact and showed no significant part deformation, however pronounced porosity was seen in the top two layers of printing. The upright triangle sample underwent significant deformation as it was embedded. Due to the base of the triangular cavity being larger than the wire diameter, some porosity was found on the underside of the wire. This experimentation led to the conclussion that a circular cavity would result in better embedding. Figure 8 shows tests that were replicated for only circular cavities. Using the optimal speed of 60 mm/min, three samples withFigure 8: Embedding of kapton coated litz wire in circular cavities362three circular cavities were printed.

Using the optimal speed of 60 mm/min, three samples withFigure 8: Embedding of kapton coated litz wire in circular cavities362three circular cavities were printed. Resulting specimen showed that there were minimal surface defects when printing was resumed over the embedded wires. From previous experience, it is expected that the small defects would be removed once additional material extrudes over the surface.4. ConclusionThe demonstration of ultrasonic wire embedding on thermoplastic substrate indicates that it is possible to fabricate 3D electronics with embedded circuitry or wire routing. Large diameter wire embedding will facilitate the fabrication process of power electronics, such as electric motor components. The presented work showed that circular cavity designs can be implemented to embed the gauge of wire necessary for power electronics.

The presented work showed that circular cavity designs can be implemented to embed the gauge of wire necessary for power electronics. The cavities not only facilitated the placement of the wire, but helped stake the wire to the substrate without significant deformation. An attempt was taken to fabricate a quarter stator of an axial flux motor, shown in Error! Reference source not found.. In this case, 28 AWG kapton coated litz wire was embedded within PC substrate to mimic the winding of a stator. Benefits of AM for this specific application compared to traditional methods, would be the cost reduction of tooling, shortening the fabrication process chain, and automation of wire winding. Although, the operational temperature and material breakdown strength are the key design criteria for an axial motor/generator, this multi process AM method can be employed after selecting the thermoplastic material that will satisfy the operational requirements.

BRID MANUFACTURING: INTEGRATING DIRECT WRITE AND STEREOLITHOGRAPHYFrank Medina*, Amit Lopes*, Asim Inamdar*, Robert Hennessey*, Jeremy Palmer**, Bart Chavez**, Don Davis**, Phil Gallegos**, Ryan Wicker* *University of Texas at El Paso, W.M. Keck Border Biomedical Manufacturing and Engineering Laboratory, Department of Mechanical and Industrial Engineering, El Paso, Texas 79968-0521 **Sandia National Laboratories, Albuquerque, NM, 87185-0958Reviewed, accepted August 23, 2005AbstractA commercial stereolithography (SL) machine was modified to integrate fluid dispensing or direct-write (DW) technology with SL in an integrated manufacturing environment for automated and efficient hybrid manufacturing of complex electrical devices, combining three- dimensional (3D) electrical circuitry with SL-manufactured parts.

The modified SL system operates similarly to a commercially available machine, although build interrupts were used to stop and start the SL build while depositing fluid using the DW system. An additional linear encoder was attached to the SL platform z-stage and used to maintain accurate part registration during the SL and DW build processes. Individual STL files were required as part of the manufacturing process plan. The DW system employed a three-axis translation mechanism that was integrated with the commercial SL machine. Registration between the SL part, SL laser and the DW nozzle was maintained through the use of 0.025-inch diameter cylindrical reference holes manufactured in the part during SL. After depositing conductive ink using DW, the SL laser was commanded to trace the profile until the ink was cured.

After depositing conductive ink using DW, the SL laser was commanded to trace the profile until the ink was cured. The current system allows for easy exchange between SL and DW in order to manufacture fully functional 3D electrical circuits and structures in a semi-automated environment. To demonstrate the manufacturing capabilities, the hybrid SL/DW setup was used to make a simple multi-layer SL part with embedded circuitry. This hybrid system is not intended to function as a commercial system, it is intended for experimental demonstration only.

This hybrid system is not intended to function as a commercial system, it is intended for experimental demonstration only. This hybrid SL/DW system has the potential for manufacturing fully functional electromechanical devices that are more compact, less expensive, and more reliable than their conventional predecessors, and work is ongoing in order to fully automate the current system.Keywords: rapid prototyping; stereolithography; direct-write; hybrid manufacturing; functional integrated layered manufacturingIntroductionRapid prototyping (RP) technologies such as solid freeform fabrication (SFF), layered manufacturing (LM), and other similar technologies enable the manufacture of complex three- dimensional (3D) parts. Stereolithography (SL) is one of the most widely used RP technologies known around the world. The resolution of SL machines and the ability of SL to manufacture highly complex 3D objects allow SL to build both functional and non-functional prototypes.

The resolution of SL machines and the ability of SL to manufacture highly complex 3D objects allow SL to build both functional and non-functional prototypes. The materials’ development for SL and other RP technologies has been primarily focused on improving the mechanical strength, durability, and thermal property characteristics of RP materials. As a result of these improvements, there is great interest in expanding the markets for RP into many areas such as rapid manufacturing of fully functional electromechanical devices.39In an attempt to assist in expanding the markets for RP into rapid manufacturing, researchers at the UTEP W.M. Keck Border Biomedical Manufacturing and Engineering Laboratory and Sandia National Laboratories have developed a hybrid manufacturing setup.

Keck Border Biomedical Manufacturing and Engineering Laboratory and Sandia National Laboratories have developed a hybrid manufacturing setup. This setup provides an integrated manufacturing environment where SL is employed in conjunction with direct write (DW) fluid dispensing technology to capitalize on each of the individual process capabilities for developing unique electromechanical devices. The hybrid SL/DW machine offers enormous potential in integrated manufacturing of unique electromechanical devices by removing many of the complex shape manufacturing limitations experienced by conventional manufacturing processes. Recent breakthroughs in DW ink dispensing for military applications have made this technology attractive for manufacturing complex 3D, multi-layered, high density, integrated, and fully functional electromechanical systems (Church et al., 2000). These electromechanical systems are more compact, less expensive, and more reliable than their conventional predecessors.

These electromechanical systems are more compact, less expensive, and more reliable than their conventional predecessors. Process development issues include improvements in STL file generation, integration of curable media such as DW conductive inks with SL technologies without contamination, adhesion of the ink and SL resin, and curing of the conductive ink. Although there have been improvements in writing capabilities, many available SL resins are incompatible with the high temperatures required to cure DW inks (typically above 100°C). Recent developments in high temperature resins such as DSM Somos® ProtoToolTM, ProtoThermTM, and NanoFormTM resins have enabled the integration of DW with SL technologies.In recent years, researchers have developed systems for automatic dispensing of conductive, thermally curable media, such as DW inks for maskless patterning of electronics.

Lewis and Gratson (2004) listed several applications of robotic dispensing systems for creating 3D arrays of many fluidic matrices. Ang et al. (2002) used a robotic dispensing system for manufacturing 3D scaffolds. Im et al. (2002) proposed the concept of functional prototype development to overcome the limitations of mono-material and mono-functions of conventional RP systems. Kataria and Rosen (2004) demonstrated a start and stop process to embed functional inserts using SL. Palmer et al. (2004) proposed a new manufacturing process to integrate SL and DW technologies for the fabrication of 3D, high-density circuitry. The work of Palmer et al. (2004) demonstrated functional electromechanical systems that are arguably more compact, less expensive, and reliable than the traditional systems by combining DW and SL in separate processes. Medina et al.

Medina et al. (2005) further demonstrated the integration of separate DW and SL systems to incorporate a video camera in a commercial clock for a home security application. Some of the build issues encountered during the work of Palmer et al. (2004) and Medina et al. (2005) include the re-registration, continuous cleaning, and curing of the part during the build process. After the SL part was built, it was removed from the SL machine, cleaned, and cured to perform the DW process. Once the DW process was completed, the part needed to be re-registered to the SL platform before continuing the build to encapsulate the circuitry (Palmer et al., 2004; Medina et al., 2005).

Once the DW process was completed, the part needed to be re-registered to the SL platform before continuing the build to encapsulate the circuitry (Palmer et al., 2004; Medina et al., 2005). These two previous efforts motivated the development of an automated manufacturing environment that combines SL and DW.thefeasibility ofrapidly manufactured,integrated,andThe current effort represents the continuing evolution of the Functional Integrated Layered Manufacturing (FILM) concept introduced in the work of Medina et al. (2005) where RP is employed in conjunction with automated access to a toolbox of additional manufacturing technologies to capitalize on individual process capabilities for developing unique functional devices. As the next step in this development, this work presents a hybrid SL/DW machine setup for rapid development of circuitry to manufacture more compact electromechanical components40in a semi-automated environment.

The hybrid system provides a method for automated deposition of fluid media during the SL build (by stopping the SL build but not removing the part from the SL machine), automated curing of the media using the SL laser and continued building without the need for recalibration and intermediate washing or resin curing. An additional linear encoder attached to the SL z-stage was used to maintain part registration during the interrupted build processes. The DW system employed a three-axis translation mechanism that was integrated with the commercial SL machine. Registration between the SL part, SL laser and the DW nozzle was maintained through the use of .025-inch reference holes manufactured in the part using SL. The current system allows for easy exchange between SL and DW in order to manufacture fully functional 3D electrical circuits and structures in a semi-automated environment.

The current system allows for easy exchange between SL and DW in order to manufacture fully functional 3D electrical circuits and structures in a semi-automated environment. To demonstrate the manufacturing capabilities, the hybrid setup was used to make a simple multi-layer SL part with embedded circuitry. The following describes the design and demonstration of the semi-automated hybrid SL/DW machine.Hybrid stereolithography/direct-write machine setupThe hybrid SL/DW manufacturing setup offers an integrated RP system comprising a traditional SL apparatus and DW fluid dispensing system as depicted in Figure 1. The SL 250/50 machine manufactured by 3D systems was used with a 355 nm solid-state laser upgrade (Series 3500-SMPS, DPSS Lasers Inc., Santa Clara, CA).

The SL 250/50 machine manufactured by 3D systems was used with a 355 nm solid-state laser upgrade (Series 3500-SMPS, DPSS Lasers Inc., Santa Clara, CA). DSM Somos® ProtoThermTM 12120 was selected as the resin for this work due to its ability to withstand temperatures (HDT) up to 126 ˚C and relatively low viscosity (550 centipoise, cps at 30 ˚C) (DSM Somos®, New Castle, Delaware). One of the issues encountered in the work of Palmer et al. (2004) was the difficult handling and building characteristics of ProtoToolTM associated with the materials high value of viscosity (2500 cps). The low material viscosity of ProtoThermTM facilitated improved handling, building without sweeping, and cleaning prior to DW deposits.UV shieldDPSS solid state laser systemDW dispensing syringe and nozzle3-axis translation stage3D Systems 250/50 SL machineDispensing system consoleTranslation stage controllerAdjustable height mobile workbenchFigure 1.

The hybrid SL/DW machine setup.41The DW setup was placed on a custom-made, adjustable height table. The table consisted of a stainless steel top and an adjustable-height hand-crank two legged mobile workbench (Part # 9054T113 and Part # 3517T82, respectively, McMaster-Carr Supply Company, Chicago, IL). Two support plates were welded to the stainless steel table top which was then bolted to the workbench base using standard fasteners. The table was redesigned to incorporate two 3/8 in. dowel pins that matched 3/8 in. holes drilled in the rim support plate of the SL machine using a hand drill as shown in Figure 2a. Alignment was achieved by inserting the dowel pins in the matching holes in the rim support plate until both pins touched the rim, illustrated in Figure 2b.

Alignment was achieved by inserting the dowel pins in the matching holes in the rim support plate until both pins touched the rim, illustrated in Figure 2b. The X-Y axis movement of the DW dispensing nozzle was provided by two automatic linear stages each 3-feet long with 40 thread/inch lead screws (Type M062-LS09, 200 steps/rev, Velmex Inc., Bloomfield, NY), arranged perpendicular to each other using support plates. An additional manual linear stage (Model # MA4012P40-S4-0, Velmex Inc., Bloomfield, NY) with an 8-inch traverse distance and 40 thread/inch lead screw was used for Z-axis movement as well as for supporting the tip of the DW ink dispensing system. The X, Y, and Z stages of the DW system facilitated movement of the fluid dispensing system, enabling precise fluid placement without intermediate calibration or removal of the SL part.

The X, Y, and Z stages of the DW system facilitated movement of the fluid dispensing system, enabling precise fluid placement without intermediate calibration or removal of the SL part. An additional z-stage linear encoder (LS 303, Heidenhain Corporation, Schaumburg, Illinois) was used to maintain part registration during build interrupts required for incorporating the DW process.Rim support plate with holes to receive dowel pinsX-Y translation stageAlignment dowel pinsZ traverse system w/dispensing tip holderAlignment dowel pins(a)(b)Figure 2. (a) The 3-axes traverse system, (b) alignment of the SL/DW system.The 2405 Ultra Dispensing System (EFD Inc., East Providence, Rhode Island) shown in Figure 3 was integrated with the SL machine and adapted to deposit fluid such as conductive ink on the part manufactured by the SL machine.

It consisted of an air pressure control console and a dispensing syringe and tip (with nozzle diameters down to 6 mils), which was connected to the console with a plastic hose. The console obtains an externally supplied inlet air pressure (up to 100 psi) and regulates the outlet pressure (from 0 – 5 psi), which was then supplied to the dispensing tip. The ink can be dispensed manually or controlled using LabVIEW® to draw given circuit profiles. It may be adapted to dispense a variety of fluids including conductive inks and virtually any other fluid. For this work, a conductive ink (CI-1002, ECM, Delaware, Ohio) was dispensed on ProtoThermTM 12120 resin (DSM Somos®, New Castle, Delaware). The ink used here was selected based on its low cost ($ 0.50/gm), low curing temperature (110 ˚C for 2 hours using 15 minute intervals), and low resistivity (0.035 Ω/square-mil) (ECM LLC, Delaware, Ohio).

The traverses as well as the dispensing system were controlled using a separate computer with LabVIEW®.42EFD 2405 Ultra Dispensing System consoleDispensing nozzleDispensing syringeFigure 3. The DW Ultra Dispensing System.A preliminary ink dispensing experiment using several dispensing tips at varying dispensing pressures aided in determining the dispensing tip and pressure selection used here and the results are shown in Figure 4. The size of the dispensing tip and the dispensing pressure were the most important parameters which determined the line width of the circuitry for a given Z distance (distance between the dispensing tip and the SL part) of 5 mils and traverse speed of 0.0625 in/sec. The selected speed was recommended by the traverse manufacturer for the type of lead screw that was used in the experimental setup. Also, this speed was selected because it was the fastest speed at which the traverse did not skip while producing a uniform ink line.

Also, this speed was selected because it was the fastest speed at which the traverse did not skip while producing a uniform ink line. The cost (up $300k) of faster (> 7 in/sec) and more precise DW systems prevented the use of those systems in the experimental setup. As stated earlier, this system is not intended to function as a commercial system, it is meant for experimental demonstration only. The dispensing tips ranged in size from 6 to 16-mils while the dispensing pressures ranged from 1 to 5 psi. The 6 mil, 10 mil, and 13 mil dispensing0.080.07) s e h c n i ( h t d W e n Lii0.060.050.040.030.020.016-mil tip10-mil tip13-mil tip16-mil tipSelected dispensing parameters00123456Dispensing Pressure (psi)Figure 4.

Line width versus dispensing pressure for different size injection tips (0.0625 in/sec traverse speed and 5-mil distance between tip and surface).43tips provided discontinuous lines below 3, 2, and 1 psi dispensing pressures, respectively (Figure 4 illustrates results for continuous lines only). The experimental results indicate that the combination of the 6-mil dispensing tip and a dispensing pressure of 4 psi provided the narrowest uniform line (0.014”), and these operating conditions were selected for the work presented here.Hybrid SL/DW machine operation overviewIn order to integrate SL and DW, the standard operation of SL was modified. Figure 5 provides a sample part with embedded electrical circuitry, vertical interconnection between the DW circuitry, and multiple SL builds that encapsulate the circuitry. This figure is intended to provide a general description of the operation of the hybrid SL/DW system.

This figure is intended to provide a general description of the operation of the hybrid SL/DW system. The figure represents building two separate SL parts (each with multiple layers), two separate DW electrical traces, and a vertical interconnection between the two circuit traces (providing 3D circuitry). The key elements of successful operation include being able to accurately register the DW trace with the SL laser (since the SL laser is used to cure the ink) and providing vertical electrical interconnections. The registration issue is solved by using a small cylindrical hole manufactured during the SL build that is used to register the DW ink dispensing tip to the SL part. The SL vat is lowered to its lowest available z-height when using DW so that the platform can remain at the same z-height. With the platform remaining at the same height, there are no geometric differences between the DW trace and the SL laser trace when curing the DW ink.

With the platform remaining at the same height, there are no geometric differences between the DW trace and the SL laser trace when curing the DW ink. To provide vertical interconnection, the reference hole (or other holes placed in the SL part) is incrementally filled with conductive ink and cured using the SL laser. These features summarize the necessary elements for successful hybrid SL/DW operation. Additional details of the operation are provided below in the description of part manufacture illustrated in Figure 5.Figure 5 represents a simple part with embedded circuitry to be manufactured using the hybrid SL/DW setup. The SL laser is used to cure the conductive ink which is dispensed during the DW process.

The SL laser is used to cure the conductive ink which is dispensed during the DW process. The laser curing process creates a continuous manufacturing scenario with theSecond DW layer profile (5th part)Hole to test 3D connectivityBuild section for encapsulating DW circuitry (3rd part)STL part to cure the ink for vertical interconnects (4th part)First DW layer profile (2nd part)Vertical interconnectThe base (1st part)Reference hole for DWFigure 5. The hybrid SL/DW operation.44potential to manufacture fully functional electromechanical components. Laser curing permits the build to continue without removing the part from the platform. The hybrid setup thus addresses the manufacturing issues like part re-registration, continuous cleaning, and curing between part builds, as mentioned in Medina et al. (2005). SolidWorks modeling software or similar CAD packages can be used to generate the individual part designs, and converted into STL file format as required.

(2005). SolidWorks modeling software or similar CAD packages can be used to generate the individual part designs, and converted into STL file format as required. 3D Systems proprietary slicing software, 3D Lightyear™ 1.4, is used to convert the STL files generated during the part preparation and design stages into the build file format required for the operation of the SL 250/50.The entire build process is designed in the form of an assembly using SolidWorks software. The assembly is first saved as a part file and then converted into an STL file. The final STL file should be divided into five distinct parts, as shown in Figure 5, to facilitate the SL/DW integration. The first STL part is the base on which the first DW circuit is drawn. This base features a reference hole which facilitates the initial calibration of the ink dispensing tip to the SL part. Traditional SL part building is performed layer by layer with planned interruptions to incorporate the DW process.

Traditional SL part building is performed layer by layer with planned interruptions to incorporate the DW process. Once the individual SL build is complete, the vat is lowered, while maintaining the build platform position constant to expose the top surface. The top surface can then be cleaned using cleaning paper dipped in alcohol. The top surface is further optionally cured using the laser to ensure no contamination between the resin and the ink. The DW system is then used to dispense the conductive ink.LabVIEW® is used to guide the 3-axis traverse and provide maximum freedom for the DW experimentation. Registration between the DW tip and the SL part and laser are achieved by locating the dispensing tip on top of the reference hole using the X-Y traverses. The initial DW starting position plays a crucial role in acquiring the required circuit profile on the SL part.

The initial DW starting position plays a crucial role in acquiring the required circuit profile on the SL part. The dispensing nozzle was controlled using the manual traverse actuator and accurately positioned over the registration hole via visual inspection (Future research will involve automating this process featuring a digital position readout). The Z-axis traverse is adjusted to 5-mils using standard slip gauges so the required gap between the dispensing tip and the SL part is obtained. After the initial calibration and accurate positioning of the dispensing tip is complete, the DW process is performed using LabVIEW®. The traversing distances and the direction of traverse are manually fed into the program.After the DW process is complete, the DPSS solid-state laser power is increased up to 100 mW by adjusting a laser attenuator located in the path of the laser beam.

The depth of resin penetration (Dp) can also be reduced from 6 to 1 mils, which reduces the speed of the laser traverse and provides increased localized heating for ink curing. The effect of curing was investigated by testing the resistivity using a MeTex® multimeter (Model # M–3850D). The vat is covered with a UV shield to protect the resin from the laser during the ink curing process. The SL build layer is selected such that the laser replicates the first DW ink profile, which is the second part of the final product design shown in Figure 5. This process is repeated until the required resistivity is attained. The ink is considered cured when the resistivity is within 3% of the previous measurement after the laser trace.The SL part is then allowed to cool to a specified temperature before the SL resin level (or vat) is raised and leveled before the next build process can begin. The next build process requires encapsulating the first DW circuitry.

The next build process requires encapsulating the first DW circuitry. The laser power is reduced to 35 mW for the SL part build and the build layer is adjusted to replicate the third part of the final product design shown in Figure 5. Once this intermediate SL build is completed, the resin level is again lowered to expose the top surface which is cleaned and cured as described above to prepare for45the next DW process. The reference hole on the third part of the SL build is filled with ink to provide a vertical connection between the two circuits. The laser power is again increased up to 100 mW to cure the ink. The hole is filled with ink one layer at a time and cured using the laser.

The laser power is again increased up to 100 mW to cure the ink. The hole is filled with ink one layer at a time and cured using the laser. These processes are continued until the desired SL/DW part is completed.Fabrication demonstrationIn order to demonstrate fabrication with the hybrid SL/DW system, a sample part was manufactured using DSM Somos® ProtoThermTM resin and the CI 1002 conductive ink according to the first 3 layers depicted in Figure 5. The SL/DW process was initiated by building the SL base with the reference hole. Once the first SL part was complete, the vat was lowered to expose the top surface, which was cleaned using paper dipped in alcohol. The top surface was further cured using the laser to prevent contamination between the resin and the ink. The 2405 Ultra Dispensing System was then aligned as described previously and used to dispense the conductive ink.

The 2405 Ultra Dispensing System was then aligned as described previously and used to dispense the conductive ink. LabVIEW® was used to guide the custom designed traverse system as well as control the dispensing for the DW experimentation.1 inch(a) Figure 6. (a) The traverse system for DW, (b) DW circuit profile(b)Figure 6a shows the hybrid SL/DW system with particular emphasis on the three-axis traverse mechanism used to guide the DW system. The SL part build was interrupted at the required layer and the DW ink dispensing system was aligned using the reference hole and commanded to deposit the circuitry as shown in Figure 6b. The integrated system was used to build circuitry without part removal or additional calibration.

The integrated system was used to build circuitry without part removal or additional calibration. The traverses were controlled using LabVIEW® to replicate the required DW profile and dispense the ink to complete the circuitry as shown in Figure 7a.After fluid deposition was complete, the ink was cured with the DPSS solid-state laser (the same laser used to perform SL) as shown in Figure 7b. Figure 8a shows an image of the cured ink. To test the effectiveness of the laser for ink curing, ink was deposited on a ProtoThermTM test piece and placed in a heat convection oven at 110 ˚C. The resistivity of the ink was monitored and recorded at 15 min intervals. The resistivity value recommended by the ink manufacturer was 0.035 Ω/square-mil.

The resistivity of the ink was monitored and recorded at 15 min intervals. The resistivity value recommended by the ink manufacturer was 0.035 Ω/square-mil. It was found that the laser curing process (10 traces at 3 minutes and 40 seconds per trace or 37 total minutes and 100 mW laser power) produced a resistivity value (0.05 Ω/square-mil), which was closer to the recommended value than could be obtained using the convection oven (0.09 Ω /square-mil) (2 hours at 110˚C). After completing the ink curing process, the integrated system leveled the vat and began traditional SL part46Laser curing the ink(a)(b)Figure 7. (a) The direct write process, (b) laser curing of the conductive ink.building on the part. The part was then reregistered to the building platform by adjusting the Z- height to allow the process to be continued at the next known layer using the z-stage linear encoder (LS 303, Heidenhain Corporation, Schaumburg, Illinois).

This was done by selecting a specific start and end layer of the build file. The remaining part was completed to encapsulate the DW circuitry using the second part of the STL file as discussed earlier. The finished part, as shown in Figure 8b, features the vertical interconnect hole and the access hole used to test 3D conductivity. The total time for manufacturing the completed part with embedded circuitry and 3D connectivity was approximately 4 hours. Work is ongoing to compare this hybrid method with traditional methods that are currently used to build electromechanical systems. The surface layer circuitry resides on the top layer of the part while the remaining circuitry is embedded within the part.Vertical interconnectAccess hole to embedded circuitrySurface layer circuitry(a) Figure 8. (a) Image of the cured ink, (b) finished part.(b)ConclusionsA hybrid SL/DW machine was developed for integrating SL and DW ink dispensing technology.

(a) Image of the cured ink, (b) finished part.(b)ConclusionsA hybrid SL/DW machine was developed for integrating SL and DW ink dispensing technology. A commercial SL machine was modified to integrate fluid dispensing or DW technology with SL in an integrated manufacturing environment for automated and efficient hybrid manufacturing of complex electrical devices. The modified SL system operates similarly to a commercially available machine, although build interrupts were used to stop and start the SL build while depositing fluid with the DW system. An additional linear encoder was connected to47the SL platform z-stage and used to maintain accurate part registration during the SL and DW build processes. The DW system featured a three-axis translation mechanism that was integrated with the commercial SL machine.

The DW system featured a three-axis translation mechanism that was integrated with the commercial SL machine. The current system allows the exchange between SL and DW in order to manufacture fully functional 3D electrical circuits and structures in a semi-automated environment.The hybrid SL/DW machine was used to manufacture a simple part featuring embedded circuitry. This work demonstrated the use of the SL laser to cure the conductive ink deposited using DW. This curing strategy enabled the build to continue without removing the part from the SL platform, maintaining accurate registration of the part within the SL machine. It was found that the laser curing process (37 minutes total laser trace time at 100 mW) produced a resistivity value (0.05 Ω/square-mil) closer to the recommended value than the convection oven curing (0.09 Ω /square-mil) (2 hours at 110˚C).

A simple strategy using reference holes manufactured during SL was successfully demonstrated for registration of the DW dispensing tip with the SL part. Vertical interconnectivity was demonstrated by filling the reference hole or other holes in the SL part with the conductive ink and using the SL laser to cure the ink. The hybrid SL/DW manufacturing setup has potential for manufacturing fully functional electromechanical components in an automated build sequence, and work is ongoing to fully automate the manufacturing process.AcknowledgmentsThe research presented here was performed at UTEP in the W.M. Keck Border Biomedical Manufacturing and Engineering Laboratory using equipment purchased through Grant Number 11804 from the W.M. Keck Foundation. This material is based in part upon work supported by the Texas Advanced Research (Advanced Technology/Technology Development and Transfer) Program under Grant Number 003661-0020-2003. Support was also provided through the Mr. and Mrs.

Support was also provided through the Mr. and Mrs. MacIntosh Murchison Chair I in Engineering and through research contract 28643 from Sandia National Laboratories in the Laboratory Directed Research and Development (LDRD) program. Sandia National Laboratories is a multi-program laboratory operated by Sandia Corporation, a Lockheed Martin Company, for the United States Department of Energy’s National Nuclear Security Administration under contract DE-AC04-94AL85000.References1. T.H. Ang, F.S.A. Sultana, D.W. Hutmacher, Y.S. Wong, J.Y.H. Fuh, X.M. Mo, H.T. Loh, E. Burdet, and S.H. Teoh, “Fabrication of 3D Chitosan–Hydroxyapatite Scaffolds using a Robotic Dispensing System,” Materials Science and Engineering, Volume C 20, 2002, pages 35–42.2. K.H. Church, C. Fore, and T. Feeley, “Commercial Applications and Review for Direct Write Technologies,” Materials Development for Direct Write Technologies, San Francisco, CA, April 24-26, Volume 624, pages 3-8, 2000.3.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperA HYBRID METHOD FOR ADDITIVE MANUFACTURING OF SILICONE STRUCTURESFarzad Liravi*, Varun Jacob-John†, Ali Toyserkani*, Mihaela Vlasea*Multi-Scale Additive Manufacturing Laboratory, Department of Mechanical and Mechatronics Engineering, University of Waterloo, 200 University Ave., Waterloo, ON, N2L 3G1, Canada† Department of Mechanical Engineering, University of British Columbia, 2329 West Mall, Vancouver, BC, V6T 1Z4, CanadaAbstractDeveloping an additive manufacturing (AM) technique for fabrication of complex-shaped silicone structures is a challenging task due to difficulties in layer-wise dispensing and stacking of such non-Newtonian viscous materials. The need for such a technique becomes apparent when surveying the ever-increasing applications of the silicone polymer in the biomedical sector.

The need for such a technique becomes apparent when surveying the ever-increasing applications of the silicone polymer in the biomedical sector. In this research, a hybrid powder-bed binder-jetting (PBBJ) and material micro-dispensing method (hybrid-PBBJ) is employed for the production of structures from silicone powder for the first time. The conventional PBBJ technique is coupled with a micro-syringe dispensing mechanism to make the layer-by-layer infiltration of fluidic silicone rubber into the silicone powder possible. Standard cylindrical (5 mm (D) ´ 3 mm (H)) and thin walled (10 mm (L) ´ 1.8 mm (H)) artefacts were manufactured as part of a design of experiment (DOE) and as a proof of concept. The AM parts were characterized in terms of geometry, porosity and mechanical performance.

The AM parts were characterized in terms of geometry, porosity and mechanical performance. The DOE results will be deployed to subsequent manufacturing of complex structures.Key wordsAdditive manufacturing, binder jetting, micro-deposition, hybrid printing, binder distribution, silicone materials1IntroductionSilicone is an organosiloxane polymer with unique properties such as high thermal stability, resistance to ultraviolet (UV) light, low surface energy, biocompatibility, and biodurability [1]. Such properties have made silicone a suitable material for many biomedical applications including prosthetics [2], medical phantoms [3], biosensors [4, 5], and drug delivery systems [6]. Silicone molding is widely used for the production of these bio-structures which imposes many restrictions on their design flexibility and functionality. The casting method is a lengthy process which may take a few days for a customized part.

The casting method is a lengthy process which may take a few days for a customized part. On the other hand, using injection molding for this purpose requires the design and fabrication of expensive casts and models that make mass-customization of such bio-structures extremely difficult [7].Additive manufacturing, as a collection of layer-based fabrication techniques, could resolve the issues associated with the current conventional methods used for the production of silicone-based biomedical products. However, AM of silicone has proved to be very challenging.1897The high viscosity of medical-grade silicones is the largest obstacle in the process of developing an AM system for the silicone, as most of the AM techniques are not designed for fluids with more than 2,000 mPa.s in viscosity.

On the other hand, the thermosetting properties of silicones prevent them from being used as a filament in the well-established fused deposition modeling (FDM) machines or in the form of resin powder in selective laser melting (SLM) methods. Furthermore, the absence of analytical models describing the behavior of viscous non-Newtonian fluids such as silicone makes the application of this material in the continuous or drop-on-demand AM systems difficult.Limited by these complications, researchers first tried to transfer the unique properties of silicone to the bio-structures made from other materials such as the commercial acrylate-based Tango Plus (Stratasys, Minnesota, USA) and the starch powder by wrapping a thin layer of silicone film around the 3D printed object [8] or dipping the object into a vat of silicone resin [9, 10].

These methods could not be established as successful bio-manufacturing techniques due to the faster failure rate of the base material compared to the silicone. As a result, developing a direct silicone AM method is crucial to overcome the mentioned problems [11-13].The most widely used AM technique for the direct additive manufacturing of silicone is material extrusion [14-20]. By deploying various types of fluid delivery mechanisms, the material extrusion technique is the only traditional AM system that can handle viscous fluids without the need for dilution [21, 22]. Freeform reversible embedding (FRE), another subcategory of material extrusion, has been introduced as a support-free solution for the 3D printing of silicone as well [23]. This technology was commercialized by the UK-based company Fripp Design and Research [24].

This technology was commercialized by the UK-based company Fripp Design and Research [24]. Although the material extrusion produces uniform high quality silicone features, it is tediously slow (5-20 mm/s) and its contact-based nature requires extensive and complex toolpath planning strategies. Photopolymerization AM such as stereolithography has been tested for the production of silicone parts without much success in terms of the geometrical accuracy [25, 26]; however, researchers are trying to optimize this process for silicone materials [7, 26]. The highest vertical resolution of ~30 µm was achieved by printing a relatively low viscous heat-curable silicone using aerosol jetting AM [27]. However, aerosol jetting is very sluggish and requires a complicated curing station to function with a high degree of repeatability.The review of AM techniques used for fabrication of silicone structures reveals that PBBJ has not been adopted for this purpose before.

This AM technique, due to its high printing speed, has the potential of being used in a commercial mass-customization production line for the silicone biomedical products. PBBJ, also known as 3D printing, was developed by MIT in 1990s [28], and involves the sequential jetting of a binder onto thin layers of powder in a powder bed to form the bulk of the structure. After printing each layer, a new layer of powder will be spread by a roller. Continuing this process will fabricate a 3D green-part supported by the loose powder which requires post-processing such as heat treatment, sintering in the furnace, or irradiation [29]. This technology has been used by multiple researchers for biomedical applications from prosthetic implants to tissue engineering [30-35].The main advantages of PBBJ systems are their fast printing speed, scalability, and the ability to print colorful structures [29, 36].

Moreover, there is typically no need for support structures in these systems, thus the design and manufacturing process is simplified, which contributes to fast production speed. In addition, PBBJ systems are capable of producing functionally graded structures [37-43], some of which are fully-colored [9, 44]. Such benefits make PBBJ an attractive technique for the fast and inexpensive development of bio-structures with1898controllable internal features. The observed limitations of commercial PBBJ systems include the relatively low resolution of internal features limited to approximately 500 µm due to difficulties in de-powdering, as well as the use of only one powder type throughout the part. Furthermore, powders are typically spread onto the build bed area in thin layers, typically 25-200 µm in thickness, which means that powders should comply with specific rheology and powder size distribution criteria for uniform and defect-free layer properties.

The parts manufactured via PBBJ are typically exposed to a post-processing protocol for de-powdering, followed by heat-annealing, chemical setting or irradiation. These post-processing steps typically result in part shrinkage. Post- processing steps reduce the geometrical accuracy of the final product. This needs to be considered in the design phase by taking into account isotropic or anisotropic compensation factors.To overcome some of the limitations in PBBJ, a new hybrid PBBJ system was developed [36, 45] to target manufacturing of functionally-graded parts with multi-material and custom variable porous structures.

This new system employs multiple powder feed mechanisms to dynamically select and deploy a variable powder composition to each layer, thermal and piezo print-head delivery systems for jetting of various liquid binders, a variable counter-rotating roller mechanism to control powder compaction, a solid material dispenser for embedding discrete pore- generating sacrificial materials (porogens), and a micro-dispensing head for deposition of liquids throughout the porous matrix of the part at select locations. A view of this system is shown in Figure 1.Figure 1 The hybrid-PBBJ AM system and examples of previously manufactured parts with variable porous properties.In this study, the hybrid PBBJ AM system is deployed for the direct production of silicone structures.

The system employs a material system composed of silicone powder as the substrate, a water-based liquid binder that is jetted through a thermal print-head to develop the layer outline of the parts, and a thermoset silicone liquid binder to impart the necessary structural integrity of each layer. The thermoset silicone binder is heat cured after printing a fixed number of layers. Curing the thermoset silicone prevents any further permeation of subsequent silicone binder through previous layers, thus preserving the dimensions of the parts. The machine and material system is used to manufacture two types of structures: a cylindrical shape and a thin walled structure as part of a DOE.

The machine and material system is used to manufacture two types of structures: a cylindrical shape and a thin walled structure as part of a DOE. The results of this work are intended to lay the ground for hybrid PBBJ manufacturing of complex-shaped silicone parts as part of a forthcoming study.18992 Materials and Methods2.1 Material System used in Additive Manufacturing of Parts2.1.1 Silicone PowderThe predominant material within the manufactured part is a spherical silicone rubber powder, coated with silicone resin (KMP-602, Shin-Etsu Chemical, Japan), which acts as a lubricant to the spherical particle, while simultaneously improving the shock resistance and spreadability. A high temperature heat resistance of the silicone powder of -50 - 250 °C allows for the silicone liquid binder to thermally set at 75 °C without damaging this silicone powder.

A high temperature heat resistance of the silicone powder of -50 - 250 °C allows for the silicone liquid binder to thermally set at 75 °C without damaging this silicone powder. The average particle size of 30 µm with a powder distribution of 4-60 µm (as shown in Figure 2) allows for the particles to closely pack together during powder spreading and binding.Figure 2 The particle size distribution of the silicone powder sample.2.1.2 Liquid silicone binderTo form the layers into a solidified silicone object, a liquid silicone rubber binder (JY- 9010, Juyou New Material Tech Co., Ltd, China), was used as-received to give the parts elastic properties through layer-wise deposition at select locations. The silicone rubber binder contains two parts, which were mixed at a 100:1 silicone precursor to cross linker ratio.

The silicone rubber binder contains two parts, which were mixed at a 100:1 silicone precursor to cross linker ratio. The silicone rubber binder is heat cured at a temperature of approximately 75 °C for 8 hours, to obtain optimum mechanical strength and ductility.2.1.3 Liquid water-based binderTo form the precursor shape of each layer, a clear water-based binder (ZB-60, 3D Systems, SC, USA) is jetted onto the silicone powder using a 600 dpi HP45 print head. The ZB-60 solution is comprised of 85-95% volume water, with the remaining volume containing humectant and a1900proprietary polymer. The surface tension of the binder is 29.05 ± 3.09 N/m. Before heat curing and evaporation, the viscosity of the binder solution is 10 Pa.s.2.2 Hybrid Manufacturing Machine SetupThe in-house developed hybrid PBBJ and micro-syringe deposition system [45, 46] was used for the production of the powder-based silicone structures.

The process consists of three main stages, performed for each layer. Firstly, the counter-clockwise rotation of the roller with the rotational velocity of 100 rpm and the lateral velocity of 20 mm/s deposits a layer of silicone powder with 100 µm thickness onto the building compartment. Secondly, the water-based ZB-60 binder is deposited into the desired shape using an inkjet printhead. The sequence of such layers was repeated a fixed number of times after which, in the third step, a liquid silicone mixture is deposited from a 100 µm nozzle through a pressurized syringe (OptimeterTM, Nordson EFD, RI, USA) to fill up the area of the desired shape, and act as the primary setting agent. The freshly deposited liquid silicone layer is then exposed to a heat lamp to cure for a short time. These steps were repeated for all the desired layers. Finally, the parts were taken out and air-heated overnight at 75 ºC to set.

These steps were repeated for all the desired layers. Finally, the parts were taken out and air-heated overnight at 75 ºC to set. Multiple components of this hybrid AM system are demonstrated in Figure 3.Figure 3 The overall view and components of the in-house developed hybrid AM system used in 3D printing of powder-based silicone structures.2.3 Manufacturing Plan2.3.1 Single droplet deposition – manufacturing of cylindrical structuresTo demonstrate the feasibility of the hybrid AM method, four cylinders (5mm D × 3mm H) were printed using the hybrid system. The printing process explained in section 2.2 was deployed, with a liquid silicone deposition frequency of every 300 µm (every 3 layers). The micro- syringe containing the liquid silicone rubber was located at the center of the cylindrical shapes 121901mm above the surface.

The micro- syringe containing the liquid silicone rubber was located at the center of the cylindrical shapes 121901mm above the surface. By applying 100 KPa air pressure, a 1.91 mm3 droplet of silicone binder was deposited on the silicone particles bound together previously by the ZB-60. After the completion of the 3D printing process, as seen in Figure 4, the parts were cured as explained in section 2.1.2 and then cleaned using a soft brush.Figure 4 The top view rendering of the build bed area after printing of the liquid silicone binder onto the silicone powder substrate.2.3.2 Single line deposition – manufacturing of thin walled structuresAs a next step, the resolution that can be achieved for a more complicated shape should be determined. To this end, rectangular shapes were made, with the silicone mixture deposited as one line down the middle of the rectangle.

To this end, rectangular shapes were made, with the silicone mixture deposited as one line down the middle of the rectangle. This is to determine the feature resolution of depositing a single line, which would be a part of producing a more complicated shape through rastering the silicone liquid deposition binder onto the substrate. To print a line in this way, the silicone was applied at the surface of the powder substrate so that it adheres to the powder upon leaving the syringe. This is instead of dispensing the silicone binder in the form of a droplet as done for the cylindrical shapes in the feasibility study described in 2.3.1. In order to see the results of the silicone printing better, black silicone pigment (Silc Pig, Smooth-on Inc., PA, USA) was added to the liquid silicone mixture so that the pigment represents 2.62% of the mixture’s total weight.For these samples, a 100 µm layer thickness was used and the powder was spread onto the designated build bed using a roller.

Then, the inkjet printhead was used to jet the ZB-60 aquaeous binder into four 10mm × 5mm rectangles, as seen in Figure 5(a). Four such rectangles were manufactured per layer. After laying down three such layers (300 µm), the syringe containing the liquid silicone mixture was lowered to the build bed until an approximately 250µm off-set from the substrate was reached. A line of the liquid silicone mixture was deposited through the middle of the previously ZB-60 printed rectangle, as seen in Figure 5(b). At this height above the powder, the liquid silicone instantly penetrated through the powder instead of forming a droplet on the nozzle of the syringe. Upon administering all four lines, the liquid silicone was immediately thermally set by using a heat lamp applied on the build bed for two minutes. These steps were repeated to create parts with a height of 1.8 mm. After the completion of the parts, the curing was performed based on the procedure explained in section 2.1.2.

These steps were repeated to create parts with a height of 1.8 mm. After the completion of the parts, the curing was performed based on the procedure explained in section 2.1.2. Parts were de-powdered after curing.1902Figure 5 The top view of the build bed; (a) After Zb60 insertion; (b) After liquid silicone deposition.2.4 Dimensional MeasurementsThe dimensions of printed walls were measured after being scanned using a 3D laser scanning microscope (VK-X 250, Keyence, Japan) at 25x magnification. Four vertical and three horizontal surface profiles were formed for each sample as shown in Figure 6.

Four vertical and three horizontal surface profiles were formed for each sample as shown in Figure 6. The average of width and height were measured over the vertical lines, and the average of length was measured over the horizontal lines.Figure 6 A sample of the dimensional measurement of the thin walled artefacts; (a) The 3D profile of a printed line; (b) The top-view composite optical-laser image of the printed line showing the four vertical and three horizontal lines; (c) A sample of the cross-section of the shape formed on a vertical line for measurement of the width and height; (d) A sample of the cross-section of the shape formed on a horizontal line for measurement of the length.19032.5 Hardness testThe hardness of the cylindrical samples was measured based on their resistance to the indentation of a shore 00 rigid ball using a handheld durometer (Shore S1, Shore Instruments- Instron, MA, USA).

The tests were conducted on the center of three cylinders with a one second dwell time, and repeated three times. Two replication was conducted for this test. A similar test was performed on a cylindrical sample with 3 mm thickness made from the liquid silicone binder.2.6 CT scanningThe internal micro-structure of the 3D printed cylindrical shapes was investigated using a nano-scale X-ray computed tomography (nanoCT) instrument (Versa 520m, Carl Zeiss AG, Switzerland) at 40kV and 727 nm voxel resolution. In order to isolate both material phases (silicone binder and silicone powder), two threshold values were manually determined as a baseline to separate the more attenuating silicone particles from the less attenuating silicone resin. To account for noise and edge hardening effects, greyscale images were first processed with both bilateral and Gaussian filters.

To account for noise and edge hardening effects, greyscale images were first processed with both bilateral and Gaussian filters. Finally, a local thickness threshold of 14.5 µm was applied to the particle segmentation to prevent brightened surface artifacts on the resin from being considered part of the particle phase.2.7Statistical AnalysisAn experimental design was formed to investigate the effects of the significant printing parameters on the dimensional accuracy of the printed line features. The goal of the study is minimizing the lateral and vertical dimension of the silicone walls while a minimum deviation between the width and height of the printed feature and the associated CAD model is maintained. DOE is a systematic method for finding the optimum treatment of input factors on the response(s). This statistical approach has been adopted by multiple researchers to improve the printing quality of different AM systems [47-50].

This statistical approach has been adopted by multiple researchers to improve the printing quality of different AM systems [47-50]. In this experiment, the parameters that can be varied are:Pressure [kPa] - The pressure applied to the syringe barrel to extrude the liquid silicone binder. Printing velocity [mm/s] - The speed of the syringe as it is depositing the liquid silicone binder. Viscosity [mPa·s] - The viscosity of the liquid silicone binder. Nozzle-substrate offset [mm] - The height of the nozzle above the powder substrate. Intermediate Curing Latency [s] - The amount of time between the deposition of the liquid silicone and the application of the heat lamp for intermediate curing. Intermediate Curing Time [s] - The time span the liquid silicone is kept under the heat lamp. Layer Thickness [µm] - The thickness of the powder rolled on at every iteration. Silicone Deposition Frequency [µm] - The stacked height of layers between depositions of the liquid silicone binder.

Layer Thickness [µm] - The thickness of the powder rolled on at every iteration. Silicone Deposition Frequency [µm] - The stacked height of layers between depositions of the liquid silicone binder. Final Curing Time [hr] - The amount of time the finished parts will be cured in the furnace.For this experiment, the effect of the first two parameters, pressure and printing velocity, was investigated, and an optimized value was selected for the other factors using trial and error method. The effect of these two parameters can be effectively quantified using a 22 factorial design experiment. The input parameters as well as their lower and upper levels are summarized in Table 1. Throughout the experiment, all the possible inputs were kept constant.

The input parameters as well as their lower and upper levels are summarized in Table 1. Throughout the experiment, all the possible inputs were kept constant. The responses monitored were the width, the height, and the length of the 3D printed features.1904Table2 2 1factorial designLevelsFactorsLow (-1)High (+1)Pressure (KPa)100200Velocity (mm/s)24In order to compare the hardness of samples printed with different treatment combinations, a one-way Tukey analysis of variance (ANOVA) at 5% significance level was conducted. Given the fact that all samples were manufactured with the same AM system and under similar conditions, they were supposed to belong to the samples with equal variances. In order to compare the hardness mean values for the 3D printed parts and the molded silicone binder, a Student’s t-test was carried out at 5% significance level.

In order to compare the hardness mean values for the 3D printed parts and the molded silicone binder, a Student’s t-test was carried out at 5% significance level. Minitab 17 software package (Minitab Inc., PA, USA) was used for all the statistical analyses.3 Results and Discussion3.1 Dimensional Accuracy of Single of the Thin-walled StructuresIn order to understand the relationship between the geometrical features of the 3D printed walls and the printing parameters, first-order empirical models were created using DOE. As explained in section 2.7, the effects of printing velocity and pressure on the width, length, and height of the walls were investigated in a 22 factorial design. The details of this factorial design as well as the values measured for each response can be seen in Table 2. Every treatment has been replicated twice in order to minimize the effect of the nuisance factors. The order of manufacturing the thin walls has been completely randomized.

Every treatment has been replicated twice in order to minimize the effect of the nuisance factors. The order of manufacturing the thin walls has been completely randomized. The analysis of variance (ANOVA) results for the width, height, and length are summarized in Table 3, Table 4, and Table 5, respectively. These results show that none of the two selected factors have a significant effect on the length and height of the walls at 95% confidence level. However, both parameters and their first-order interaction significantly control the width of the features with the velocity having the highest impact. The Pareto chart for the width (Figure 7) confirms these findings. Also, the residual plots for the width shown in Figure 8 do not show a significant deviation from the underlying assumption of experimental design.

The width of an extruded line for a low viscous polymer can be related to the pressure and velocity according to Eq. 1 [45, 51]. Base on this model, the width of the line is1907directly proportional to the pressure and inversely proportional to the velocity. A similar conclusion could be made using the regression model fit to the width using DOE (Eq. 2), with pressure having a positive regression coefficient and the velocity a negative coefficient. The accordance of the regression model to the physical rules governing the flow of fluids in the extrusion-based systems along with the high R-squared and adjusted R-squared values (97% and 96%, respectively) for the DOE validate the empirical model. The path of optimization demonstrated in Figure 9 shows that decreasing the pressure value to 100 KPa and increasing the velocity to 4 mm/s results in the line features with the highest resolution.

The path of optimization demonstrated in Figure 9 shows that decreasing the pressure value to 100 KPa and increasing the velocity to 4 mm/s results in the line features with the highest resolution. The image of this optimized thin wall as well as its 3D profile is demonstrated in Figure 10.𝑊𝑖𝑑𝑡ℎ =𝜋𝑅)𝑃 8𝜇𝑉ℎ𝑊𝑖𝑑𝑡ℎ = 2861.6 + 433.3×𝑃 − 541.7×𝑉 − 434.9×𝑃×𝑉where 𝑅 is the internal radius of the nozzle, 𝑃 is the pressure, 𝜇 is the viscosity, ℎ is the estimated height of the line feature (the nozzle off-set from the substrate), 𝑉 is the printing velocity.Curing the silicone binder immediately after dispensing creates a solid barrier, impeding the subsequent infiltration of the liquid silicone binders from the layers above. This solid barrier is the reason behind the constant height of the thin walls for different parameter combinations. The insignificant effect of the input parameters on the length of the features can be traced back to the effect of the overfilling, i.e.

The insignificant effect of the input parameters on the length of the features can be traced back to the effect of the overfilling, i.e. depositing more fluid at the start and stop points of the raster, giving a dumbbell shape to the linear features. Also, the perpendicular orientation of the wall length with respect to the roller’s moving direction results in the transferring of the wet powder along the width of the features and not their length. These observations will be more investigated and verified through future experiments.The optimized resolution achieved in this statistical analysis is approximately 2 mm which is less refined than the standard resolution required for additive manufacturing. The width of the printed line features need to be decreased so that this hybrid system can be used for the production of complex 3D structures. In the future, similar thin walled artefacts will be manufactured using a mixture of the silicone powder and polyvinyl alcohol (PVA).

In the future, similar thin walled artefacts will be manufactured using a mixture of the silicone powder and polyvinyl alcohol (PVA). Authors believe that adding PVA to the materials system will improve the geometrical accuracy of the system by preventing the excessive lateral permeation of the silicone resin. A complete thermal analysis of the silicone binder would also help in this regard by determination of the appropriate curing latency and intermediate curing time to prevent lateral spread.

A complete thermal analysis of the silicone binder would also help in this regard by determination of the appropriate curing latency and intermediate curing time to prevent lateral spread. Moreover, the current pneumatic micro-syringe system will be replaced by a solenoid drop-on-demand extrusion-system to make a better control of the deposition volume and frequency possible.1908(1)(2)4.03.52500 3000 3500W < 2500 3000 – 3500 – 4000 – 4000 >y t i c o l e V3.02.52.0100120140160180200Pressure0004W00032 0000022V3 yticole4001051PerusserFigure 9 2D and 3D contour plots showing the path of optimization; the low level of pressure and high level of velocity yield the optimized resolution.Figure 10 The optimized thin walled artefact and its 3D profile manufactured at 100 KPa pressure and 4 mm/s velocity.

View of actual part and laser-scanned artefact.19093.2 Cylindrical Parts Mechanical PerformanceTable 6 summarizes the hardness values measured for both cylindrical additively manufactured parts and the molded and cured silicone binder. The Tukey ANOVA study was carried out to investigate the deviation of samples’ averages from the population average value. The following hypothesis was formed:§ Null hypothesis: All average values are equal. § Alternative hypothesis: At least one average is different.The ANOVA test results (Table 7) do not show any statistically significant difference between the average hardness values which represents a uniform manufacturing condition over the building area (P-value: 0.086). The manufacturing conditions have also been kept constant between consecutive printings as the two manufacturing batches have equal average values at 95% confidence level.

Moreover, the simultaneous 95% confidence intervals for the average difference of all the samples include zero which confirms the findings of ANOVA.Table 6 The durometry results for the 3D printed cylinders and molded silicone binder.Hardness (shore 00)Sampletest 1test 2test 3AverageCylinder 1 (batch 1)80.585.680.982.33Cylinder 2 (batch 1)87.385.282.885.10Cylinder 3 (batch 1)80.780.181.580.77Cylinder 1 (batch 2)81.585.38684.27Cylinder 2 (batch 2)85.384.385.284.93Cylinder 3 (batch 2)80.987.884.784.47Total average for cylindrical samples83.64Molded silicone binder 181.580.175.479.00Molded silicone binder 281.281.279.680.67Total average for molded silicone binder79.83Table 7 ANOVA results for the hardness test.SourceDFAdj SSAdj MSF-ValueP-ValueMeasurements555.4911.0982.540.086Error1252.474.373Total17107.961910Figure 11 Tukey simultaneous 95% confidence intervals; CX-Y (X: Cylinder number, Y: Batch number).In order to compare the hardness of the 3D printed structures and the silicone binder, their average values were compared at a 95% confidence interval using a t-test with the following hypotheses:§ Null hypothesis: The two averages are equal.

§ Alternative hypothesis: The two averages are not equal.Based on the t-tests results summarized in Table 8, we have failed to reject the null hypothesis. Hence, the two averages are not statistically different at 95% confidence level (P-value: 0.073). The estimate of the difference between the hardness of the 3D printed structure and the pure silicone binder is +3.81 which is negligible. This statistic means that the hardness of the 3D printed sample is mostly contributed by the liquid silicone binder. The hardness of the silicone powder reported by the manufacturer is approximately 70 (shore 00). The distribution of cured silicone binder in the silicone powder at multiple levels can be seen in Figure 12.

The distribution of cured silicone binder in the silicone powder at multiple levels can be seen in Figure 12. The interconnected channels of the silicone binder throughout the structure have given the final 3D printed part a hardness value closer to that of the cured molded silicone binder.A full mechanical characterization should be carried out on the products of this AM system once it is capable of fabricating more complex objects.

Specifically, the effect of the printing parameters on the tensile strength, tear strength, and elongation at break of the parts [52] should be investigated in order to make any required changes to the chemistry and composition of the system of materials for each particular biomedical application.Table 8 t-test results for comparing the hardness of 3D printed structures and molded silicone binder.NMeanStDevSE Mean3D printed structures683.641.720.7Molded silicone binder279.841.180.831911Figure 12 The CT scan results at 727 nm voxel size; (a) The cylindrical shape and its 3D profile; (b) The 3D view of a small portion of the 3D printed structure; (c) a top view cross-section of the part near the surface; (d) a top view cross-section of the part at the center of the structure.

(c-d) The blue, yellow, and black areas correspond to the cured silicone binder, silicone powder, and air respectively.4 ConclusionIn this study, a novel hybrid PBBJ and micro-dispensing technique for the production of structures made from silicone powder and liquid silicone rubber at a proof-of-concept stage was introduced. Even though a sole PBBJ system could be much faster than the hybrid system, binding the silicone powder particles using the micro-dispense technique is necessary because current thermal and piezoelectric jetting systems are not compatible with high viscous silicone binders, specifically with thermosets. The method was successfully employed for the manufacturing of simple geometries such as cylinders and thin walls. The manufactured parts demonstrated enough resilience and hardness (approximately 80 shore 00) to be used for different biomedical applications. However, more mechanical characterization is required for future applications.

However, more mechanical characterization is required for future applications. Furthermore, the porous structure achieved can be altered for the fabrication of drug-delivery parts with controlled drug release rates.1912The designed experiment resulted in a validated regression model for controlling the width of the printed line features. The current lateral resolution of more than 2 mm also revealed that a thermal analysis of the silicone binder as well as modifications to the powder system and the AM set up are required to control the dimensions of the final products and achieve the geometrically accurate parts. These changes should be made in the future so that complex structures can be printed using this hybrid AM technique.

These changes should be made in the future so that complex structures can be printed using this hybrid AM technique. To take advantage of all the benefits of the PBBJ system, trials for developing a system based on the direct jetting of silicone binder should be continued at the same time.AcknowledgmentThe authors appreciate the funding support received from The Natural Sciences and Engineering Research Council of Canada (NSERC), grant #50503-10713. Furthermore, the authors would like to acknowledge the work of research interns Abhinav Grover and Alysan Young in developing printing recipes for the silicone material systems.5 References[1] Noll W. Chemistry and technology of silicones: Elsevier. 2012[2] Butler DF, Gion GG, Rapini RP. Silicone auricular prosthesis. Journal of the American Academy of Dermatology 2000; 43: 687-690[3] In E, Walker E, Naguib H. Novel Development of 3D printable UV-curable Silicone for Multimodal Imaging Phantom.

A review of materials, fabrication methods, and strategies used to enhance bone regeneration in engineered bone tissues. Journal of biomedical materials research Part B: applied biomaterials 2008; 85: 573-582[38] Hutmacher DW, Sittinger M, Risbud MV. Scaffold-based tissue engineering: rationale for computer-aided design and solid free-form fabrication systems. Trends in biotechnology 2004; 22: 354-362[39] Hutmacher DW. Scaffolds in tissue engineering bone and cartilage. Biomaterials 2000; 21: 2529-2543[40] Castilho M, Dias M, Gbureck U, Groll J, Fernandes P, Pires I, Gouveia B, Rodrigues J, Vorndran E. Fabrication of computationally designed scaffolds by low temperature 3D printing. Biofabrication 2013; 5: 035012[41] Hollister SJ. Porous scaffold design for tissue engineering. Nature materials 2005; 4: 518- 524[42] Szucs TD, Brabazon D.

BRID PROTOTYPES TO ASSIST MODELING AUTOMOTIVE SEATS J. C. Ferreira*, H. Madureira† and R. Beira† * Mechanical Engineering Department, Instituto Superior Técnico, 1049-001 Lisboa, Portugal † IN+, Design Engineering, Instituto Superior Técnico, 1049-001 Lisboa, PortugalAbstractThe development of new modular seats is an important issue in the automotive industry. However, is very time consuming and costly. Virtual models and hybrid prototypes could accelerate the car seats development process. The hybrid prototypes are mainly manufactured by rapid prototyping with multi materials. The objective of this paper is to establish a methodology to develop innovative lightweight multi-functional, modular car seats to be used in Multi-Purpose Vehicles (MPV), by means of FEA simulation and rapid prototyping additive/subtractive technologies utilizing multi materials. A case study is presented to validate the developed methodology.

A case study is presented to validate the developed methodology. The manufactured hybrid prototype’s reproduces the main functionalities of the MPV modular seat, namely its three key positions: normal, stored and table.IntroductionToday manufacturers experience immense pressure to provide a greater variety of complex products in shorter product development cycles. The evolution of the market needs the time-to-market reduction, mainly because the product life cycle is shorter and also because it is very important to produce more rapidly from an initial conception or “idea” to a mass production product. New approaches in rapid product development (RPD) from the design point of view are needed [1]. The late discover of product requirements are fairly common and the use of certain methods and product representations or materializations may reduce this problem.

The late discover of product requirements are fairly common and the use of certain methods and product representations or materializations may reduce this problem. The identification and implementation of designer requirements in the early stages of RPD are significant issues for a successful product development reducing time [2]. These needs, coupled with the desire to further reduce costs and improve quality have impelled manufacturers to turn their focus toward integration of product design with faster manufacturing process activities. One possible methodology to achieve this goal is through the use of rapid prototyping (RP) technologies giving the ability to transit quickly from concept to rapid solid freeform fabrication reducing costs and improving product quality faster. RP represents technologies that quickly realize the conceptualization of a product design.

RP represents technologies that quickly realize the conceptualization of a product design. The evolution of RP allows obtaining parts representative of mass production products within a very short time.The main advantages of virtual prototyping (VP) in the RPD context consist in the ability to alter earlier the shape at the designer’s will, directly in CAD format. The design visualization or virtual prototyping (VP) can also be carry out handling stereolithography tessellation language (STL) models, the de facto standard for layer manufacturing technology (LMT) [3]. The VP in the RPD framework allows finite element analysis (FEA) simulation techniques of intermediate geometric models to optimize the design prior to fabricating the physical model. Careful examination of the visually simulated model before the actual fabrication can help minimizing unwanted design iterations [4].

Careful examination of the visually simulated model before the actual fabrication can help minimizing unwanted design iterations [4]. From the improved VP can be manufactured conceptual models by rapid prototyping (RP) in a short time (or mask time overnight) to evaluate the set up corrections. RP provides true CAD design verification, because digital data is used directly to manufacture a cost effective prototype.25The basis of rapid prototyping technology allows the rapid SFF of models built by additive and/or subtractive RP [5]. The rapid production of a prototype make available manufacturers having a better product development overall control, not only of the product itself - visual aids for engineering, ergonomics, and fit - but also of the processing technology by having a prototype at an early stage of the rapid manufacturing process chain.

RP responds better to the growing interest of the industry in reducing the time to market of new products and respective costs [6].The RP in wax material can be obtained directly by multi-jet modelling (MJM) techniques. The surface finish and dimensional precision of parts obtained through this rapid prototyping process is often highly important, especially in those situations where the manufactured prototypes are used as functional lost patterns, as in the investment casting technology [7]. The RP wax pattern dimensions are determined by the wax's thermophysical and thermomechanical properties and RP process parameters [8].The principal bridge between RP and SFF is based on the indirect use of STL geometry data files for additive and/or subtractive manufacturing.

These tessellated representations of products have several drawbacks linked to the shape and topology of the surface triangular mesh.The virtual prototyping simulation and analysis in the RPD methodology contest provides the ability to adjust earlier the geometric shapes to fulfill the design requirements prior to fabricating the corresponding solid freeform physical parts [9].RPPD MethodologyThe purpose of this R&D work was to investigate the crucial roles of RP technologies and purpose a practical methodology to integrate advanced rapid technologies in simultaneous engineering for rapid product and process development (RPPD).

The developed methodology provides meaningful information for manufacturers to integrate VP/RP/RE/SFF technologies that are best suited to their production requirements in order to reduce the product development leading time and reducing manufacturing costs.The rapid product and process development (RPPD) methodology for rapid productdevelopment integrating VP/RP/RE/SFF is schematized in Fig. 1.The RPPD methodology starts with the product specifications from designers and the 3D- CAD design serves as the starting point in the integration of product and process development allowing starting the production in a shorter time.

Once the 3D-CAD design has been analyzed and the corrections completed, it serves as a link to utile rapid prototyping for solid freeform fabrication process.In this study, the RPPD methodology linked with several technologies includes: CAD software to generate file data models based in functionality and ergonomics (*.stl format); virtual prototyping for preliminary FEA stress analysis; SFF technologies to built conceptual prototypes; simultaneously the metrology control is perform aided by reverse engineering; rapid SFF of physical hybrid prototypes by composite technology, followed by surface finishing; functional tests with the physical hybrid prototypes.

After some iteration to improve the product26performance and enhance the manufacturing process in a short lead time, they are approved and the mass production could start.Figure 1 – Methodology for Rapid Product & Process Development.The actual know-how in RP allows saying that it is important to utilize 3D-CAD software to define parts with solid modeling capabilities in the preliminary design stage avoiding either wire-frame or surface models. A solid freeform model enables cross sectioning to expose internal details and computation of moment of inertia. The solid freeform model also allows the designer to easily determine part parameters such as weight and centre of gravity through the software. These parameters are crucial to the design evaluation of a part. The 3D-CAD and VP software serves as a powerful “tool” in design evaluation, particularly in the areas of finite element analysis (FEA) and mechanism evaluation design by virtual prototyping animating the mechanisms.

FEA predicts stress and deflection on a structure or allows process simulation.A solid freeform model of a part can be imported from a CAD system in the STL file format to a RP system. The STL file approximates the surface of the solid freeform model by covering it with a multitude of triangles. Subsequently, a RP system, which uses the technology such as fused deposition modeling, builds a part with the SFF process that grows the part in successive additive layers. Ideally, to make a good RP model, the STL file must be made from a watertight solid model. If the user manages to completely seal and close all corners and seams, any of the major CAD software packages is capable of providing true solid freeform models and27consistently exporting flawless STL files. Otherwise, corrections can be made in dedicated software’s.

Otherwise, corrections can be made in dedicated software’s. Designers can make decisions faster and iterate design changes more quickly using prototypes, saving time and energy that translates into saving product and process development costs. With a 3D model in hand, designers can easily identify costly design flaws or some manufacturing unfeasibility. As a result, modifications of a part are made early in the design process before costs escalate. Detecting design flaws with prototypes complements evaluating design alternatives with engineering based CAD software. Also the prototype accuracy controlled fast by metrology aided by scanning and reverse engineering (RE) indicates eventual geometry deviations.The use of SFF systems enhances the integration of design.

It reduces the number of costly changes as it enhances communication between the design and manufacturing departments, giving manufacturing a chance to integrate manufacturability expertise and influence the product design while changes are inexpensive. The cost of a change increases roughly by an order of magnitude as the design proceeds from the significant development phase to the next. The other advantage when using rapid SFF to make functional hybrid prototypes is that components may be manufactured with low and high resistance materials (e.g. expandable wax or polymeric materials assembled with aluminum or steel articulation joints). Some SFF functional prototypes can even be cut, drilled, tapped, sanded and finished in a number of ways for higher quality.

Some SFF functional prototypes can even be cut, drilled, tapped, sanded and finished in a number of ways for higher quality. At this point, the non-standard functional prototypes (SFF devices) can be utilized mounted as components together on testing apparatus for functional and ergonomic studies.Numerical Modeling of the MPV SeatThe module Design aims developing a lightweight multi-functional, modular and innovative automotive seat, to be used in a MPV-type (MPV - Multi-Purpose Vehicle) City Car. The seat’s external shape has been defined by designers and then draft in CATIA 3D-CAD software. The virtual prototype has proved very helpful in resolving conflicts between the seats main three moving parts, but at a given stage it was clearly insufficient as a working tool for engineering analysis.

Consequently, was decided to build a semi-functional 1:5 scale prototype to serve as reference for discussions and also as a marketing tool for presentations to the community.The numerical modeling of a light weight MPV seat aimed at giving realistic boundary conditions to the human model. The modular MPV seat consists in various mechanical elements. The main parts of the seat were: the head-rest clamp, the head-rest foam, the foam of the backrest, the foam of seat base, the backrest spring, and the cover of the seat. The geometry of the MPV seat was design based on new ideas and is illustrated in Fig. 2a.To model the solid freeform modular MPV seat the experimental data used in this studywere completed in a context of ergonomics and comfort [10 – 11].The designer’s team modeled the external freeform of the MPV seat using Rhinoceros 3D modeling software.

However, this virtual model was meant to convey only the general impression of the shape and so it was composed of unconnected and nonintersecting surfaces that proved to be extremely difficult to “mend”. Consequently a different approach was followed, which consist in importing into CATIA the Rhino model, using the IGES import/export standard format [12].28The exact surfaces (not class “A”, but at least closed and generally continuous) were then modeled using the designers’ freeform as a reference. A general overview of the process is represented in Fig. 2b.(a)Figure 2 – a) Representation of the MPV modular seat modeled in Rhinoceros 3D software b) The CATIA modeling process starting from designer freeform.Simulation and FEA Analysis with MPV ModelIn this study, the material behavior laws for the foam and the cover were considered linear with material properties resulting from the literature. The seat backrest and cushion were more detailed than the seat base.

The seat backrest and cushion were more detailed than the seat base. Thus, a simplifying hypothesis consisted to model seat backrest and cushion with flexible shell elements which aimed at limiting the movement of the human body.The mechanical properties of base seat was been extracted from compression tests and was considered with linear rigid shell elements and the Young modulus was of 1000MPa and the Poisson's ratio was 0.3. The thickness of the shell elements was of 3.0mm with a density of 7.9 Kg/l.Special attention was paid to the backrest and cushion frames of the seat. The backrest and cushion frames were modeled with shell elements. The frame geometry was simplified. It was divided into three parts: the base frame considered as rigid body; the seat cushion frame which can be deformable; and the seat backrest frame also considered as a deformable body part.

It was divided into three parts: the base frame considered as rigid body; the seat cushion frame which can be deformable; and the seat backrest frame also considered as a deformable body part. The normal position of the seat cushion and the backrest frame were related by a fixed articulation as a rigid body.The material properties of the backrest and cushion steel frames were determined in order to have a qualitatively realistic behavior under static loading. We thus obtained a Young’s modulus of 235MPa and a Poisson's ratio of 0.28. The density was of 550.0 g/l and the thickness of the frames was of 1.75mm.

We thus obtained a Young’s modulus of 235MPa and a Poisson's ratio of 0.28. The density was of 550.0 g/l and the thickness of the frames was of 1.75mm. The stiffness of the mechanism between the back rest and the cushion frame were chosen very high to be considered as rigid.Car Seat FEA Model: The first 2D analysis from simulation of the homologation test ECE-R 14 on two different light-weight seat structures, using ABACUS software was performed29(b)only to get some feeling for the structures ability to resist the kind of loads involved, Fig. 3a. The preliminary 3D static nonlinear analysis has proved that the backrest and the base frames of the MPV axisymmetric light-weight seat structure was too weak to support the loads involved, Fig. 3b.(a)(b)Figure 3 – a) 2D simulation on a new MPV seat structure with ABACUS software. b) 3D static nonlinear analysis of the new MPV seat structure.The backrest was reinforced, as shown at 3D static nonlinear analysis in Fig.

b) 3D static nonlinear analysis of the new MPV seat structure.The backrest was reinforced, as shown at 3D static nonlinear analysis in Fig. 4a Also the seat base was reinforced, as shown at 3D static nonlinear analysis in Fig. 4b. In terms of general resistance, the results resemble the real life test well enough.(a)(b)Figure 4 – a) 3D static nonlinear analysis of reinforced backrest frame. b) 3D static nonlinear analysis of reinforced seat base.30Hybrid propotypes by SFF processPreparation of Rapid Prototyping Process: In order to export a valid STL file, one must have a completely close surface. Although CATIA offers a number of tools to verify the quality of surfaces, the existence of gaps and lacks in continuity, and the ability to export a surface in STL format, experience has shown that the risk of incomplete or “wrong” prints was strong when trying to print directly from surfaces.

An alternative approach was then taken, which consisted in first converting the CATIA surfaces in a solid freeform, Fig. 5, and then generating the STL file.Figure 5 – Modeled SFF virtual prototype of the MPV seat and a detail.3D Printing the Prototype: After the STL file was ready it was imported into 3D Systems’ [13] proprietary software, which allows a number of operations to be carried out before starting the 3D printing process.The most important of these operations is part placement. Since material is added to support the prototype build up its position has a strong influence on where the support ribs are fixed on the surfaces. Derived from the need to build supports for surfaces with inclinations greater than 2-3 degrees with respect to the vertical there are two types of surfaces in the final model: high quality surfaces where there were no supports and low quality rough surfaces in the places where supports had to be built.

This is important, since the necessary removal of those support ribs can damage the prototype surfaces, or at least roughen these surfaces. Generally the surfaces that are vertical or facing upwards are good quality, whereas surfaces which normal is down facing are rough. The need for good quality surfaces in specific zones of the model determines the model placement within the build volume.The functionalities that the software includes are: estimation of the total build time; STL file verification; previewing of the supports to be built and automatic part placement. Another factor that is influenced by part placement is build time.

Another factor that is influenced by part placement is build time. The 3D Systems’ ThermoJet takes approximately six minutes to build every millimeter in height, so by carefully examining the part31orientation and other requirements one can achieve the best compromise between surfaces quality and build time.Hybrid Seat Prototype SFF by 3D Printing: In this work a ThermoJetTM solid object printer from 3D Systems, using the multi-jet modeling (MJM) technique is used to produce solid freeform physical models of designed MPV seats. This system uses a print head comprising 352 jets oriented in a linear array. The large number of jets allows fast and continuous material deposition at a resolution of 300 x 400 x 600 dpi (xyz), producing models with a maximum size of 250 x 190 x 200 mm (xyz) and with a layer thickness of 0.042 mm [14].Making the 3D prototypes with the rapid prototyping machine involves the conversion of STL files of the parts in the machine’s software.

To produce the solid freeform physical model of the part, the corresponding STL file is transformed in machine code through the ThermoJetTM printer client software, powerful software that enables to verify the STL file, auto-fix errors and determine the better position of the model on the working apparatus area, optimizing both build space and time.Solid freeform 3D printing is done layer by layer with “Thermojet 88” wax material from3D Systems (which properties are in table 1), over the apparatus metallic platform.Melt temperature [ K ]358 -368Softening temperature [ K ] Density @ 413 K [ g/cm3 ] Density @ 403 K [ g/cm3 ] Density @ 296 K [ g/cm3 ]3430.8460.8480.975Volumetric shrinkage from413 K to room temperature [ % ]12.9Ash content (Gray wax) [ % ]0.00 – 0.01Table1: Properties of wax material “Thermojet 88” (3D SystemsTM [13])As the machine selected is a concept modeler from 3D Systems that has limits to the dimensions of the parts that could be printed.

For that reason, was necessary to rescale the car seat geometry to a smaller scale 1:5.The build process took approximately 3 hours to complete, after which it was necessary to cool the model down in a refrigerator for about 1 hour to facilitate the supports removal that broken easily.After all the RP parts had been “cleaned” they were ready for the next step of the process: the custom fitting of sheet aluminum reinforcements in the zones that support rotation axles and on the lower part of the model, where the forces exerted during model manipulation would exceed the resistance of the thermopolymer. Upon assembling the SFF hybrid prototype reproduced all the main functionalities of the MPV modular seat, as shown in Fig. 6.32Figure.

Upon assembling the SFF hybrid prototype reproduced all the main functionalities of the MPV modular seat, as shown in Fig. 6.32Figure. 6 - SFF hybrid prototype (scale 1:5) made by RP multi-jet in wax reinforced with aluminum at the zones that support rotation axles and on the lower part of the model.HSM and Hand Finishing MPV Pure Seat Hybrid Prototype: The MPV pure seat hybrid prototype, scale 1:1, was built out of several PUR blocks that were HSM (High Speed Machined) according to the 3D-CAD geometry, Fig. 7. MDF plates were used to model the components that required extra resistance. The rotations axles are made out of steel.Fig. 7 - The solid freeform hybrid prototype (scale 1:1) performs accurately the main movements of the MPV seat: (a) normal; (b) stored; and (c) table.33The solid freeform hybrid prototype has been fabricated by the industrial design team, respecting the placement and dimensions of the internal components.

This solid freeform hybrid prototype was reverse engineered to update the original 3D-CAD geometry with the changes performed on it.ConclusionsThe present paper describes a methodology for the development of efficient modular seat models, applicable for usage in the development process of light-weight modular MPV seats. Additionally, hybrid prototypes made by SFF providing appropriate input for the seat models have been described. Hybrid prototypes are mainly manufactured by rapid prototyping with multi materials. As an example, the methodology has been evaluated for one modular MPV car seat: a virtual model of this seat has been created containing the mechanical properties requested for a real seat. The following can be concluded:A methodology has been outlined to develop innovative lightweight multi-functional, modular car seats to be used in Multi-Purpose Vehicles (MPV). Virtual models and hybrid prototypes have allowed speed up the car seats development process.

Virtual models and hybrid prototypes have allowed speed up the car seats development process. This methodology allows realistic predictions of seat design and interactions with e.g. human models.The development of modular seat models by usage of numerically FEA simulation techniques for seat modeling is an efficient method in terms of computational time, for the development of seats. This strongly affirms the proposition that seat models are attractive for parameter variation at an early stage of the design process.

This strongly affirms the proposition that seat models are attractive for parameter variation at an early stage of the design process. The presented methodology is attractive for simulations with long CPU times that do not require a detailed analysis of deformations.The SFF of hybrid prototype’s with additive/subtractive techniques utilizing multi- materials reproduces the main functionalities of the MPV modular seat, namely its three key positions: normal, stored and table.In a further development it will be possible to perform a parametric study on seat characteristics and optimize the seat against the biomechanical response of the human body complex.AcknowledgementsThe authors would like to thank the entities that have supported this research: ACECIA,ACE, ModusDesign, FMH and PIEP.References 1. Bernard A., Fischer A., “New trends in rapid product development”, CIRP Annals-Manufacturing Technology, vol. 51 (2), 2002, pp.635-652.2.

Hybrid Rapid Pattern Manufacturing System for Sand Castings Matthew C. Frank, Frank E. Peters, Xiaoming Luo, Fanqi Meng, Joseph Petrzelka Department of Industrial and Manufacturing Systems Engineering Iowa State University, Ames, IA, 50011, USAAbstract This paper presents a Rapid Pattern Manufacturing system developed for the sand casting process. It involves both additive and subtractive techniques whereby slabs are sequentially bonded and milled using layered toolpaths. As such, patterns are grown in a bottom-up fashion, both eliminating the need for multi-axis operations and allowing small features in deep cavities. Similar approaches exist in the literature; however, this system is specifically targeted at large wood and urethane sand casting patterns. This method introduces a novel support structure approach by integrating a flask into the pattern build process.

This method introduces a novel support structure approach by integrating a flask into the pattern build process. We also present adaptive slicing algorithms that optimally place layer transitions to avoid thin sections near flats, peaks, and valleys or where interaction with chemically bonded sand could be problematic. The system has been implemented in an automated machine capable of producing patterns in excess of several thousand pounds. Preliminary testing of the system in the development of next generation military equipment is presented in a case study.Introduction Sand casting is utilized in the manufacturing of a wide range of metal part sizes, from one to several thousand pounds. Even though sand casting has been used for centuries, it is still one of the most important manufacturing processes today. A key element in the sand casting process is the pattern used to form the mold cavity in sand.

A key element in the sand casting process is the pattern used to form the mold cavity in sand. Once a pattern is made, tens, hundreds or sometimes thousands of molds can be made; each producing a part. There is a strong motivation to use a rapid prototyping technology for pattern manufacturing, especially for short runs or prototyping where the costs of a pattern cannot be easily justified. However, there are limitations in the current RP technologies mostly related to size constraints and materials available. Most sand casting patterns are made from wood, although some are made of urethanes and metals. Pattern making is considered a highly skilled task and most patterns today are made by specialty pattern shops that serve the foundries, although some foundries still maintain pattern making departments. In early times, patterns were made manually by craftsmen using lathes, mills and other woodworking machines.

In early times, patterns were made manually by craftsmen using lathes, mills and other woodworking machines. In some cases, the pattern shop not only creates, but also designs the pattern geometry given the desired part geometry. For example, the original designs of the parts need to be modified to determine parting lines, design cores, apply shrink rules and add draft to surfaces for the subsequent pattern geometry. The emergence of modern CNC machines has reduced the need for hand- made or manually processed patterns; however, this has only shifted the requirements of the pattern makers to high-skilled NC programmers and machine operators. Using a CNC router or milling machine provides the necessary geometric and material capabilities for the pattern industry; yet a truly automated or Rapid technology is still not available. Rapid Prototyping & Manufacturing (RP&M) techniques emerged only a few decades ago.

Rapid Prototyping & Manufacturing (RP&M) techniques emerged only a few decades ago. Early adopters of some technologies were pattern making shops that needed a better method for testing part and/or pattern designs. This allowed different shrink, draft and gating systems to be tested by making a few sand molds and pouring metal. Once the design was finalized, a durable pattern could be manufactured using conventional means.Related Work An early technology was Laminated Object Manufacturing (LOM), which could create sand casting patterns that, at least in appearance and size, were very close to patterns used in the foundry. LOM was used for sand casting patterns, wax injection molds for investment casting, and master models for silicone molding (Mueller and Kochan, 1999). Wang et al.

LOM was used for sand casting patterns, wax injection molds for investment casting, and master models for silicone molding (Mueller and Kochan, 1999). Wang et al. (1999) discussed the LOM process for sand casting patterns and concluded that LOM-based rapid tooling yielded about a 50% time and cost savings compared to aluminum tooling, but that some geometry may not be suitable and that errors in the pattern were common. Selective Laser Sintering (SLS) has been used to directly create the sand molds using coated (croning) sand. This technique can be used for very low-volume sand casting molds with complexgeometries which would be difficult to create using traditional sand casting mold techniques (Tang et al., 2003). Similar to SLS, Stereolithography (SLA) has been explored as an important technique in the rapid tooling field.

Similar to SLS, Stereolithography (SLA) has been explored as an important technique in the rapid tooling field. The SLA investment casting build structure called QuickCastTM was introduced in 1993 and has been used to create functional parts in a variety of different metals (Jacobs, 1995; Hague et al., 2001). SLA parts have also been used as patterns to prepare RTV molds and epoxy molds for injection molding (DirectAIM) (Karapatis et al., 1997). Fused Deposition Modeling (FDM) has been directly or indirectly used in investment casting; for investment casting using wax FDM parts or to produce RTV molds from FDM plastic parts to create wax investment casting patterns (Lee et al., 2004). Three Dimensional Printing (3DP) has also been used to create sand casting patterns and molds. The Zcast technique from ZCorp was developed in order to print sand casting molds directly (Kawola, 2008).

The Zcast technique from ZCorp was developed in order to print sand casting molds directly (Kawola, 2008). Thermojet printing such as the Sanders ModelMaker has been effectively used for investment casting through the printing of wax. Its small layer thickness enables very smooth part surfaces, but it also increases the manufacturing time greatly and has relegated the technology to small parts such as jewelry (Naitove, 1996). There has been some research in other technologies that use subtractive methods and hybrid approaches using additive and subtractive means. Schaaf (2000) presented a sand mold RP technique using industrial robots. Yang et al. (2002) presented a Robotic Machining RP system using a 6-axis robot on a linear track to perform the cutting operation and a rotary platform to position the workpiece. Hur et al. (2002) created a hybrid system using machining and deposition.

Hur et al. (2002) created a hybrid system using machining and deposition. Shape Deposition Manufacturing (SDM) decomposes the CAD models into sections that can be deposited as near-net shapes and then machines each to net- shape before depositing and shaping additional material (Merz et al., 1994).Solution Methodology A proposed methodology for Rapid Pattern Manufacturing this is presented (RPM) section. The basic premise is to utilize a combined additive and subtractive approach in order to take advantage of accepted materials used in pattern making, while enabling very simplified toolpath and process planning so that it can be automated. The concept is to stack “slabs” of material which are subsequently machined to specified “layer” heights and with the required 3D geometry of only that layer. As such, very deep and complex patterns can be machined using considerably short tools. The process planning is reduced to a set of 2- and 2½-D toolpaths for each layer.

As such, very deep and complex patterns can be machined using considerably short tools. The process planning is reduced to a set of 2- and 2½-D toolpaths for each layer. For each layer, a face mill reduces the slab to a layer height and then a sequence of flat- and ball-milling operations using waterline toolpaths generate the 3D surface geometry. The basic steps of the process are illustrated in Figure 1.L1S1Mill to layer (L1) w/geometryMill to layer (L2) w/geometryinAdd slab (S1)Add slab (S2)Figure 1 – Basic steps in the Additive/Subtractive approachMachining each slab to a particular layer thickness serves dual purposes; 1) it creates an accurate flat surface to ensure proper bonding of the subsequent layer and control over the part height and 2) it allows control over where the “seams” occur along the build height. The Additive/Subtractive Rapid Pattern Manufacturing system presented in this paper is designed for 3-axis, single-sided milling.

The Additive/Subtractive Rapid Pattern Manufacturing system presented in this paper is designed for 3-axis, single-sided milling. As such, the process is suited mainly for the creation of two-part patterns for the metal casting industry. As partly an academic exercise, we include one undercut geometry (flat surface) for layer placement consideration, since they can be effectively created using an appropriate placement of slabs and layers. However, there are practical instances to justify this, since riser elements (used in pattern design for properly filling a casting) are sometimes implemented as loose components that may have undercut features. In these cases considering undercut flats in the algorithm has practical, albeit not straightforward uses.The advantage of this rapid pattern manufacturing method may not be obvious, as many in the industry currently use a similar approach with much larger “slabs”, whereby blocks of material are glued together and then machined altogether.

The advantage to our proposed method is two-fold 1) we reduce the process planning step from a large machining process plan to simpler, individual layers and 2) mostimportant, one can feasibly machine an entire pattern with very deep cavities and small features (illustrated in the example of Figure 2. Using this additive/subtractive approach, we can always use short, small diameter tools as needed, regardless of depth. Collision conditions typically avoided using 5-axis systems are eliminated altogether in our approach. As shown in Figure 2a, a large slab or solid block approach will lead to inaccessible regions. In contrast, our approach allows a small tool to access small regions, since the subsequent (higher) layers have not yet been added. In the laboratory, we have demonstrated this by making as small as 3mm features (interior radii) close to 1 meter deep in a pattern cavity.

In the laboratory, we have demonstrated this by making as small as 3mm features (interior radii) close to 1 meter deep in a pattern cavity. A sample pattern for a military component is illustrated in the implementation section; a smaller pattern, with only ~1meter x-y by 0.5meters deep. In theory, the system has very few limits on feature size, since we can control layer depth; hence, the maximum length tool required (short tools can have small diameters as needed, long tools cannot).(a)(b)Figure 2 - Deep cavity example (a) large slab or solid block approach causes collision, (b) Layer based approachThe specific problem addressed in this paper is choosing the thickness of the layers comprising each build, based on the slab thickness and the part geometry. As in many additive RP systems, a uniform layer thickness could be used, yet this may result in in-process failures and/or problems in the final quality of the pattern.

As in many additive RP systems, a uniform layer thickness could be used, yet this may result in in-process failures and/or problems in the final quality of the pattern. These machining conditions have resulted in fractured, chipped and/or rough surfaces on wood patterns, or exposed adhesives at layer interfaces. Laboratory experiments have shown that the variable layer placement improves the process significantly; resulting in the successful processing of complex patterns with good surface finishes. This problem of layer placement in RP is not new, however, it is typically not motivated by the same processing requirements. Previous researchers have studied layer thickness in additive/subtractive manufacturing, such as Hur, et al (2002) who presented a hybrid rapid prototyping system using machining and deposition based on a STEP feature model.

In a layer based robot machining system presented by Chen and Song (1999), layer thickness is determined based on the feature visibility and slab thickness. Binnard and Cutkosky (1998) utilized a pre-defined basic shape library to facilitate layer thickness planning for SDM. Pinilla, et al (1998) presented another layer thickness method in SDM which was based on the analysis of all silhouette edges that denote transitions from non-undercut surfaces to undercut features. Chang, et al (1999) presented a layer thickness planning approach based on surface splitting. The Free Form Thick Layered Object Manufacturing (FF-TLOM) is a technology that enables the fabrication of large shapes from thick layers of foam with smooth non- facetted surfaces (Broek et al., 2002). The Solvent Welding Freeform Fabrication (SWIFT) process repeats the cycle of solvent welding and CNC contour machining on material sheets (Cormier and Taylor, 2001; Taylor et al., 2001).

The Solvent Welding Freeform Fabrication (SWIFT) process repeats the cycle of solvent welding and CNC contour machining on material sheets (Cormier and Taylor, 2001; Taylor et al., 2001). The uniform stock layer thickness in SWIFT is limited by the feeding system, which introduces geometric error (Yang et al., 2002). Song, et al (2005) presented a direct approach for freeform fabrication of metallic prototypes by 3D welding and milling. Adaptive slicing (Sabourin et al., 1997; Tyberg and Bohn, 1998) also deals with the layer thickness problem; however, the layer thickness definition in the adaptive slicing is different from the layer thickness in this paper. The objective of layer thickness decisions in adaptive slicing is to enable contours in each slice to best represent the part geometry in an efficient manner.

The objective of layer thickness decisions in adaptive slicing is to enable contours in each slice to best represent the part geometry in an efficient manner. However, the layer thickness decision in the proposed Rapid Pattern Manufacturing system is to ensure part geometry is machined effectively, given the geometry of the pattern and the tools and materials used to create the pattern. In previous work, most researchers have considered layer thickness with a motivation of part geometry realization (to make it possible to create the geometry), while some have also considered the material slab thickness constraint. In the proposed Additive/Subtractive Rapid Pattern Manufacturing system, geometry realization is not a problem in theory; two-part patterns for casting components with a definable parting line is not a problem.

In contrast, this work is motivated by in-process failures and the final surface quality and strength of the pattern, which we believe can be significantly affected by layer thickness/layout. The problem is to develop an algorithm that will take as input the surface geometry of the desired part (pattern/mold/etc.)and determine an effective sequential strategy for applying slabs and creating layer thicknesses. The general solution methodology involves several key areas of investigation including; 1) determining a set of factors affecting layer thickness decisions, 2) evaluating the input geometry to determine important “features” of the geometry, 3) conducting a feature height analysis, and 4) determining layer thicknesses appropriate for each unique combination of feature heights. It should be noted that although the term “feature” is referred to in this paper, the goal of the research is to provide a more or less “feature-free” input requirement.

It should be noted that although the term “feature” is referred to in this paper, the goal of the research is to provide a more or less “feature-free” input requirement. That is, typical feature-based approaches assume that a part model is pre-defined by a set of “features” such as holes, planes, slots, cavities, bosses, etc. Therefore, when this paper refers to features, it is in the sense of geometric characteristics of the part geometry, rather than traditional constructive solid geometry.n oPlane Ii t c e r i dFactors affecting layer thickness The main factors affecting the layer thickness decision in this Rapid Pattern Manufacturing are based on a few assumptions about the general system setup.

In this work, the following process and setup is proposed: 1) thick slabs of material are stacked and bonded on a build platform 2) a 3-axis CNC mill/router machines each slab to a flat layer of designated height and forms the part surfaces within that layer, and 3) a set of cutting tools is available, with lengths as long as the slab is thick, or the maximum layer thickness, as required. This paper proposes that the layer thickness criteria are then based on 5 factors; (1) Minimum cutting tool length,(2) Material slab thickness,(3)Part geometries, (4) Slab and bonding strength, and (5) Freeform surface slope. The minimum cutting tool length from the set of available tools determines the cutting depth for the system; therefore the maximum layer thickness is constrained by this value. The layer thickness must obviously be less than or equal to the material slab thickness.

The layer thickness must obviously be less than or equal to the material slab thickness. With respect to part geometries, as shown in Figure 3, Plane I can be created; however, Plane II can only be created if a layer transition occurs precisely at this height. This is only possible when the plane is parallel to the faces of the slab (perpendicular to the stacking direction). Any other down-facing features, such as Plane III, cannot be fabricated by this system. When machining, cutting forces can be sufficient to damage a very thin layer, regardless of the bonding strength of the adhesive, or the thin section may vibrate if bonding is not complete. In addition, it is undesirable to have the adhesive exposed as a large surface on the part. In practice, these are potential places where chemically bonded sand could stick to the pattern.

In addition, it is undesirable to have the adhesive exposed as a large surface on the part. In practice, these are potential places where chemically bonded sand could stick to the pattern. These reasons make it necessary to have a minimum criterion for the thickness of each layer.Plane IIIdl iPlane IIu bFigure 3 – Basic part geometriesFeature Analysis: A “feature” in this paper is loosely defined as a portion of a part having some machining significance and can be fabricated using 3-axis single-sided machining/routing. According to this definition, there are obviously countless surface shapes that could be considered features. In order to simplify the problem, these features are divided into 3 major groups, Type I, Type II and Type III features. Type I features include local peaks, local valleys and up-facing flats. Type II features are limited to planes having a normal in the –z direction. Finally, a Type III feature is a freeform surface with a shallow slope.

Type II features are limited to planes having a normal in the –z direction. Finally, a Type III feature is a freeform surface with a shallow slope. As illustrated in Figure 4, a local valley exists as the bottom of a slot (Figure 4c) while local peaks exist on the top of the spherical and rounded entities (Figures 4a,c). The up-facing and down-facing flats are simply flat surfaces with normals in the +z or –z direction, as illustrated in Figures 4b,d. The position of a Type I feature is directly related to the thin layer problem described in the previous section. These heights along the z-axis must be found such that layer transitions at(a)Local Peak (Type I)(c)Local Valley (Type I)(e)(b)Up-facing flat (Type I)(d)Down-facing flat (Type II)Small Slope surface (Type III)Figure 4 – Feature Examplesthese heights are avoided.

In contrast, Type II features (Figure 4d) dictate precisely where a layer transition must occur, since it is impossible to create these undercuts by 3-axis machining along the z- axis. Therefore, the bottom side of the slab actually becomes the down-facing flat of the Type II feature. Of course, this also implies that the surface accuracy of the Type II entities is dependant on the slab surfaces, since they will not be machined surfaces. The reader will note that the Type II feature (a down facing plane) is included in this system because it is possible to create them; however, if the system is used for a purpose such as a sand casting pattern Type II features will not exist because they cannot release the sand mold, unless they are a special case of a loose piece riser that will be cut separately. A Type III feature (shallow sloping surface), affects the machining quality similar to an upfacing flat in that we can be confronted with thin material conditions and/or exposed adhesive.

A Type III feature (shallow sloping surface), affects the machining quality similar to an upfacing flat in that we can be confronted with thin material conditions and/or exposed adhesive. The feature heights are the basis of the layer thickness algorithm, therefore the first step is to calculate these heights. The following section presents methods to determine the location of Type I, Type II and Type III feature heights. The location of the these feature heights along the z-direction will be used in conjunction with the slab thickness and layer parameters in order to determine the most effective layout of layer transitions. The goal in layer positioning will be to avoid these critical features.Type I: Local peaks and valleys: The local peak or valley presents a problem of thin materials in convex or concave surfaces. The local valleys have the potential to expose a considerable amount of adhesive.

The local valleys have the potential to expose a considerable amount of adhesive. In practice, this has posed a problem if the resins used in the chemically bonded molding sand for casting react with the adhesive. The problem with peaks can be more catastrophic, as we have experienced material failure during cutting. As shown in Figure 5a when an arbitrary layer placement leaves a small contact patch for the next layer, the machining of the slab may shear off the feature. Granted, the adhesive bond does not typically fail, in fact the cyanoacrylate glues used in the process are stronger than the MDF material; hence, the MDF material fails. Examples of these catastrophic failures are presented in the implementation section.(a)(b)Figure 5 - Local peak issueFor this analysis, we simply analyze the slice geometry from an STL file.

Examples of these catastrophic failures are presented in the implementation section.(a)(b)Figure 5 - Local peak issueFor this analysis, we simply analyze the slice geometry from an STL file. Each slice of an STL file contains several loops, or polygonal chains, and each chain defines part of the cross sectional slice of the object at that given layer height. When a loop appears or disappears from one slice to a successive slice of an STL file, it indicates the emergence or disappearance of what is referred to as a feature in this research. The feature heights can be obtained by locating these emerging and disappearing loops within the cross sectional slices of the part geometry as the slices are searched along the z-direction. There are several slicing algorithms available (Luo et al., 2001; Choi and Kwok, 2002; Pandey et al., 2003), thus it is easy to obtain the loops from a STL file, and then feature heights can be acquired by comparing these 2D loops.

In previous work, Tyberg (1998) presents a contour vertical connectivity matching method. The method computes the intersection of two contours which belong to the same sub- slab. However, this approach becomes computationally expensive if many contours exist in each slice. In contrast, we will use a two-step method to speed up the local peak and valley search process. If the numbers of loops in two adjacent slices are different, there must be a feature appearing and/or disappearing. Of course, if the number of loops in these two adjacent slices are the same, it does not necessarily follow that there are no feature changes between them. For example, if an equal number of features appear and disappear simultaneously, then the total loop count for each slice will be the same. Therefore, the first step is to check the numbers of loops in these two adjacent slices. If numbers are different, a disconnection is detected.

Therefore, the first step is to check the numbers of loops in these two adjacent slices. If numbers are different, a disconnection is detected. If the loop numbers are equal a containment evaluation across the slices is performed to assess if a feature is present. This containment relationship analysis is based on a Point Containment Assumption, as follows: If two points having the same coordinate value in the x-y plane are located separately in two planes bounded by line loops on adjacent cross sectional slices, then these two loops are assumed to be from the same part body. However there is an exception in reverse: As shown in Figure 6, planes S and S′ are bounded by line loops from two different part bodies, and there areS⊂ . This exception exists only when the distance separating loops two points p and p′ with p is less than the resolution of the slicing algorithm.

This exception exists only when the distance separating loops two points p and p′ with p is less than the resolution of the slicing algorithm. One simple, but costly method to solve this problem is to use an extremely small slice spacing; however, that could be computationally expensive for tall parts. Of course, the precision of feature detection for our algorithm is decided by the slice resolution; if slice spacing is large, feature height precision suffers. In order to quickly and accurately locate the feature heights, a Halving Algorithm (Matthews and Fink, 2004) is adopted. In this manner, a relatively low resolution slice spacing (~ 0.1 inch, 2.54mm) can be used to initially search for features, and then a smaller resolution (~ 0.01 inch, 0.254mm) is used to precisely locate feature heights. In reality, the probability of having the same amount of features disappear and appear at the same z height is very small.

In reality, the probability of having the same amount of features disappear and appear at the same z height is very small. Therefore, the probability for this connectivity detection algorithm to get to step 2 is small, making the method more computationally efficient.S⊂ ,p''Figure 6 – Exception to point containment assumptionType I: Up-facing flats: The halving is process appropriate for determining heights of local peaks and valleys; however, it is inefficient and inaccurate in finding the exact height of a horizontal plane. In the case of up-facing flat features, it is known that there must be some facet with its normal parallel to the +z axis direction (has a (0,0,1) normal vector, as illustrated in Figure 7a. Therefore, the facet normals of the STL file are searched to locate these potential heights of up-facing flats. In related work, Emmanuel (1996) searched for continuous groups of triangle facets which share the same z height to detect horizontal areas.

In related work, Emmanuel (1996) searched for continuous groups of triangle facets which share the same z height to detect horizontal areas. It should be noted, however, that some small triangular facets created to fill holes in the STL model (post-process repair algorithms), may have the same normals, but do not necessarily represent a flat planar feature (Figure 7b). Another issue is that a small up-facing facet could occur at the tangent “peak” of a freeform or otherwise curved feature. A method to filter these instances is as follows: 1) if two or more adjacent triangles have +z normal, then they exist on an up-facing flat feature (avoids detecting peaks) and 2) if only one triangle whose normal is in +z direction is found, and one dimension of the triangle is significantly small (smaller than the chordal deviation of the STL model), this triangle is not part of up-facing flat feature (avoids triangles added via repair programs).

For example, vertex P on the triangle in Figure 7c is very close to the edge L, because this is a very thin triangle added during STL generation/repair. Obviously, this cutoff value can vary depending on the scale of the model and chordal deviation, but it should be a straightforward parameter to establish. Once all local peaks and valleys and up-facing features are determined, their heights are stored into what will be called Data Set I. This data set helps determine candidate locations for layers to exist throughout the build height. Although they (peaks, valleys and planes) are located in one data set, they are not treated equally, depending on how close together they exist along the build height. In the case where a local peak or local valley height is within a default distance to an up-facing plane, the local peak or local valley feature height is deleted from Data Set I.

In the case where a local peak or local valley height is within a default distance to an up-facing plane, the local peak or local valley feature height is deleted from Data Set I. This approach is employed because an up-facing flat feature is more critical for layer placement as it generally has larger surface area (at the designated z-height), compared to a local peak or valley, hence it is more important to avoid a large area of exposed bonding material as a pattern surface.(a)(b)(c)NdescribedaboveLNPZYXFigure 7 – Detecting up-facing flat; (a) up facing flat, (b) small up- facing facet not on a plane, (c) parameters to define small facetsType II: Down facing flats: Both the down- and up-facing flats have normals parallel to the z axis, albeit in opposite direction.

Therefore the Type II feature height analysis method is the same as the method for up-facing flat height analysis described above; but leads to placing down-facing flat heights into a second set called Data Set II. As opposed to the up-facing flat heights in Data Set I, for each down-facing flat inData Set II, there must be a new layer at that height in order for the down-facing flat to be generated by the bottom face of the material slab.Type III: Shallow Sloping Freeform Surfaces: The STL file format approximates freeform surfaces with many triangular facets; hence, the idea of analyzing slopes of surfaces is reduced to simply analyzing the triangles of the STL file, and all points in each triangle facet have the same slope. The motivation to study this freeform surface shape is similar to avoiding up facing flat, as a shallow surface approaches the same characteristic. As shown in Figure 8, an arbitrary layout may result in very thin material.

As shown in Figure 8, an arbitrary layout may result in very thin material. Assuming some level of incomplete adhesion (as stated previously, experienced in the laboratory), these thin regions are potentially sheared off during cutting. It should be noted that this problem is similar to the local peak condition illustrated above (Figure 5), where a thin section fractures under cutting forces. The angle of the facets to the horizontal is found and we establish a minimum value experimentally/experientially based on the strength of the pattern material. In laboratory experiments a nominal value of 15° has successfully avoided chipping in Medium Density Fiberboard (MDF), a material that is suitable for sand casting patterns.

In laboratory experiments a nominal value of 15° has successfully avoided chipping in Medium Density Fiberboard (MDF), a material that is suitable for sand casting patterns. Higher strength materials such as hardwoods, RenBoard, or of course metals would allow smaller shallow facet angles.Thin regions(a)(b)(c)New methodOne subtle difference between Type III features and the other two features discussed above is that a Type III feature often covers a range along the Z direction since the freeform surfaces are approximated by many small triangles in the STL file format, rather than a distinct height of a peak, valley or flat. The total range of Type III features along the Z direction are determined by the range of z-heights for the vertices of all shallow slope facets and these feature ranges are stored in Data Set III. It should be noted that material slab heights are limited therefore; a Type III feature may not always be avoided in layer height calculations.

It should be noted that material slab heights are limited therefore; a Type III feature may not always be avoided in layer height calculations. In this case, the methods of the secondary approach presented later in this paper are utilized.Figure 8 - Shallow slope surfacesLayer thickness algorithm The proposed layer thickness algorithm defines locations where slabs of material are bonded in the additive portion of the process. After a slab is placed and bonded onto the stack, the subtractive process not only creates the 3D geometry of the layer, but also mills the slab to the designated layer thickness. Although the slabs are typically of uniform thickness, layer thicknesses will vary throughout the part as required. Slab thickness could in fact also be varied, if for no other reason than to reduce waste (reduce the amount of material removed when the layer height is much smaller than the slab thickness).

Slab thickness could in fact also be varied, if for no other reason than to reduce waste (reduce the amount of material removed when the layer height is much smaller than the slab thickness). For each layer, the slab could simply be chosen as the smallest slab that is thicker than the current layer thickness. This small improvement is ignored in this paper, as it does not change the layer thickness algorithm development. Moreover, allowing a variety of slab thicknesses adds considerable complexity to the Rapid Pattern Manufacturing System; one would need to be able to store, pickup and place a variety of thicknesses. In the current system, we have only used a uniform slab thickness based on available pattern materials in sheet form (i.e. ~0.75” MDF boards).Primary layer thickness strategy for Type I features: Firstly, we establish a minimum thickness allowable for any layer, denoted MTmin. This is a default value dependent on the material strength and bonding strength.

This is a default value dependent on the material strength and bonding strength. The maximum layer thickness LTmax is set to the minimum value between the slab thickness ST and tool length TL; LTmax = Min (ST, TL), where MTmin is the minimum material thickness, ST is the material slab thickness and TL is the minimum tool length. The primary layer thickness strategy for each Type I feature is: There should be no deposition within a z region (H1m - MTmin, H1m) which iscalled a Non-Deposition region, where: H1m is the mth Type I feature z height. To begin, Non-Deposition are calculated and stored in data set NDj, which is the Non-Deposition z height region (NDj-, NDj+). Each NDj includes two values: the upper limit NDj+, and the lower limit NDj-; where NDj- = H1m – MTmin and NDj+ = H1m ( j = m ). If two + ).

Each NDj includes two values: the upper limit NDj+, and the lower limit NDj-; where NDj- = H1m – MTmin and NDj+ = H1m ( j = m ). If two + ). If there exist Non-Deposition regions overlap, they are combined as NDj- = ND(j-1)- ( If NDj - < ND(j -1) Type II entities in the range of a Non-Deposition region, the Non-Deposition region is then divided into two Non-Deposition regions at these Type II entity z heights, since a Type II feature has higher priority. This division operation avoids layer placement conflicts between Type II entities and Non-Deposition regions. The sample part illustrated in Figure 9 has seven Type I feature entities; hence there are seven Non-Deposition regions. It happens that some Non-Deposition regions in this example connect; hence only two Non-Deposition regions are formed after combining. Secondary layer thickness strategy for Type I features: After all combinations, some Non-Deposition regions may exceed the height of the material slab thickness.

Secondary layer thickness strategy for Type I features: After all combinations, some Non-Deposition regions may exceed the height of the material slab thickness. This indicates that it not possible to have a layer covering the entire Non- Deposition region. When this problem occurs, the primary layer thickness strategy for Type I feature fails, and a secondary layer thickness strategy is employed. When the height of a Type I feature is above the layer position by MTmin, bonding strength and material strength are assumed sufficient to ensure proper machining and part quality. However if the distance between the feature and the layer position is less than MTmin, the quality/value of a layer height choice is decided by two factors: 1) the distance between the feature and the layer position (DT), and 2) the cross section loop area (A) at the particular layer z height.

When two Type I feature entities connect to each other, the best position to place a layer along this Non-Deposition region is exactly at the lower entity position, which has the largest DT and A for the upper entity. Therefore, the lower feature positions are evaluated using a simple cost objective function to quantify the benefits of placing layers at these positions; Q = α * DT + β * A, where: α is a material coefficient (unit: (inch)-1 ), β is a bonding/adhesive coefficient (unit: (inch2)-1), DT is the distance between feature and layer position, and A is the cross section loop area at the layer height. The material coefficient α and bonding/adhesive coefficient β are weight parameters which specify the influence of DT and A on the Q-value. Currently, these two coefficient values were acquired through experimental efforts with medium density fiberboard (MDF) as the pattern material and Cyanoacrylate as the adhesive.

Currently, these two coefficient values were acquired through experimental efforts with medium density fiberboard (MDF) as the pattern material and Cyanoacrylate as the adhesive. For general implementation, there is no theoretical optimal weight values that would have a straightforward derivation. Rather, they would need to be tuned experimentally based on the pattern material and adhesives of choice (stronger or weaker materials or glues could dictate arbitrarily different weights). This heuristic approach does not necessarily yield an optimal solution based on Q-value, especially since it may result in an excessive number of layers in a Non-Deposition region, driving up material and adhesive costs.regionsFigure 9 – Non-Deposition region combinationThen, a branch-and-bound algorithm was adopted to minimize the number of layers, while maximizing the Q-value of possible positions in the Non-Deposition regions.

The first level selects the number of layers placed in the Non-Deposition region. The search starts from the theoretical minimum number of layers, which is Lmin = UR [ ( NDj+ - NDj- ) / LTmax ], where UR(x) is an operation that calculates the smallest integer ≥ x. Next, feature heights are searched in sequence of increasing Q-values for possible layer placement solution. Consider the ND2 in Figure 9 for example. The height of ND2 is 1.0 inch, LTmax is 0.8 inch, f6 (which has the largest Q-value) is 0.1 inch from the bottom of ND2, and f5 has the second largest Q-value, which is 0.25 inch from the bottom of ND2. For level 1, the minimum layer number needed to cover the ND2 in theory is UR(1.0/0.8) = 2. Then, the level 2 starts from f6. If f6 = 1, at least one more layer is required since UR((1.0-0.1)/0.8) =2, and the total number of layers is greater than 3.Therefore, f6 = 0.

Then, the level 2 starts from f6. If f6 = 1, at least one more layer is required since UR((1.0-0.1)/0.8) =2, and the total number of layers is greater than 3.Therefore, f6 = 0. In level 3, if f5 = 1, no more layers are required, since UR((1.0-0.25)/0.8)=1, and the total number of layers required is 2 which is equal to the minimum theoretical number of layers. At this point, the search process is completed and layer positions in the Non-Deposition region are stored into Data Set II, which stores the Type II feature heights. Layer placement in the Non-Deposition region is actually similar to the Type II feature problem, in which there must be layers placed at these positions to cover Non-Deposition regions.Overall Layer Thickness Algorithm: The overall layer thickness algorithm places layers in a bottom-up fashion.

To begin, the bottom of the CAD model of the pattern is positioned at z = 0, and the initial start position is set to Hc = 0, where Hc is the current tentative layer z height. The current layer number i is set to 1, H10 is set to 0. When the search begins, the current tentative layer height is set to one maximum layer thickness (LTmax) higher than the previous layer: Hc = Hi-1 + LTmax. Next we determine Layer Thicknesses for Data Set II. Type II features and layer positions in Non-Deposition regions are evaluated first since layers must correspond exactly at theses heights in order to create the down-facing flat geometry, or to provide suitable machining quality in the Non-Deposition regions. First, all height data in Data Set II are searched to determine those that are within the height range Hi-1 to Hc.

First, all height data in Data Set II are searched to determine those that are within the height range Hi-1 to Hc. If multiple heights are possible, the height that is closest to Hi-1 is the next layer position (Hi), where Hi is the z height of the ith layer If no such feature height is found, the search continues for Type I feature heights. For Layer +,then Hc is in the thickness for Type I features, if there is a j that meets the condition: NDj- < Hc < NDj range of a Non-Deposition region. As such, Hc should be moved out of the Non-Deposition region Hc = NDj-; else, a layer can be placed at Hc directly. The Non-Deposition regions here are only those less than LTmax. If they are greater than LTmax, layers are placed directly in the first step. Therefore, Non- Deposition regions above LTmax cannot be selected in the second step. Finally, we determine Layer thicknesses for Type III features.

Therefore, Non- Deposition regions above LTmax cannot be selected in the second step. Finally, we determine Layer thicknesses for Type III features. For this step, Layers are also placed based on Type III features; however, it occurs that Type I and Type III features may cause conflicts; one cannot always satisfy both Type I and III features simultaneously. Again from experimental tests, Type III features create poor machining conditions, but Type I features more often cause catastrophic failures: hence, Type I features are given higher priority to be satisfied. We will define H3o as the oth Type III feature z height region + (that is, this Hc is (H3o- , H3o located within the range of a Type III feature), then, the height H3o- is tested. If no j exists for: NDj- < + , then Hc is moved to H3o-, or, Hc is kept the same. In this iterative manner, the Type I, Type H3o- < NDj II and Type III feature searching processes determine the layer thicknesses for the entire part.+).

In this iterative manner, the Type I, Type H3o- < NDj II and Type III feature searching processes determine the layer thicknesses for the entire part.+). For the Hc from step (b), if there is exists an o for H3o- < Hc < H3oImplementation The layer thickness algorithm has been implemented and several patterns have been created in the laboratory, with chemically bonded sand molds pulled from the patterns and casting performed. The evaluation of the approach presented in this section involves 1) comparing the calculated feature heights from the algorithm with design feature heights, 2) fabricating sample test patterns to evaluate the efficacy of this layer based approach, and 3) practical testing of the methods in the creation of actual sand casting patterns for a relatively large casting.Test Sample: The thickness algorithm was implemented in C++ on a Pentium 3.0GHz PC running Windows XP.

The input to the layer thickness software was an STL file (ASCII, 0.001 inch chordal deviation). A sample part was designed to verify the layer thickness algorithm such that all steps and conditions would be tested. In this example, the material slab thickness (Medium Density Fiber board) was 0.70 inch. Tool lengths are larger than the slab thickness, so LTmax is set to 0.70 inch. The minimum layer thickness MTmin was setlayerFigure 10 – Sample part 3D modelto 0.20 inch. The material coefficient was 0.7 (Length unit: inch), the bonding coefficient was set to 0.3 (Area unit: inch2) and the small slope surface threshold angle was set to 15º. Figure 10 shows a 3D model of the sample part. Twelve machining features are detected by the layer thickness software, with the positions of these twelve features listed in Table 1.Table 1 – Design vs.

Twelve machining features are detected by the layer thickness software, with the positions of these twelve features listed in Table 1.Table 1 – Design vs. Detected feature heights (Unit: inch)Detected HeightDesign HeightDetected HeightFeatureFeatureDesign HeightUp-facing flat 10.30.3Local peak 10.6530.65The small differences between the design positions and detected positions come from two sources. One error source the approximation inherent with an STL model while the other is from the Halving Algorithm which can only acquire approximate local peak or valley positions. That being said, in this example the differences are less than 0.005 inch and consequently have little influence on the layer thickness evaluation. Layer thickness results are presented in Figure 11. The first layer thickness follows the primary layer thickness strategy for a Type I feature; and is 0.2inch lower than local peak 1.

Layer thickness results are presented in Figure 11. The first layer thickness follows the primary layer thickness strategy for a Type I feature; and is 0.2inch lower than local peak 1. The second layer thickness utilizes the secondary layer thickness approach for a Type I feature, and the layer thickness calculated from local valley 1 meets the optimization condition (4) presented in section 2.4.2. The third layer thickness is obtained directly from the Type II feature’s down-facing flat 1. The last layer ends at the top of the part (an up-facing plane).

The system is comprised of 4 major functional elements including; 1) two elevator platforms serving as feed and build chambers with 1.2m3 (1440kg) capacities, 2) a material handling system to clamp, position and compress up to 1.2m2 sheets of material, 3) a glue application system, and 4) one off-the-shelf component; a 3-Axis CNC router. A total of 7 controllable axes are utilized in the completely automated processing of patterns. The gluing system utilizes a peristaltic pump which directs cyanoacrylate adhesive through a manifold applicator head. The servo driven build table with 4 ball screws can position the pattern for cutting operations and apply up to 17,000N of force during the 30second gluing compression cycle. The thickness algorithm has been implemented in software as a C-hook in the MasterCAM CAD/CAM environment.

The thickness algorithm has been implemented in software as a C-hook in the MasterCAM CAD/CAM environment. NC code for each layer and the requisite slab sequencing and facing to layer height data is output from MasterCAM and then processed using customized control system software to drive the machine elements.layer116cmFigure 12 - Rapid Pattern Machine in Rapid Manufacturing and Prototyping Lab at ISUThe system has been utilized to create numerous prototype patterns and most recently for a pattern of a steel cast component over 800x800x300mm. This large sand casting pattern made by the Rapid Pattern Manufacturing machine and the sand mold created from this pattern are shown in Figure 13. Figure 13a, shows the pattern in the latter stages of being machined in the system, illustrating the advantage of using a layer based approach, as it can be seen how the current layer in the picture is breaking through to reveal the deep pattern cavity below.

One will note that this pattern is considerably large and deep; however, only a 1 inch (25mm) long end mill was required, since each layer is machined after being stacked. As such, we were able to use as small as a ¼”(6mm) diameter ball mill to access small corner radii, even in the deepest regions of the pattern. Total machining time for this size pattern is currently at approximately 50 hours, or roughly 2 hours per inch (25mm) of z-height on this 116x116cm (slab dimension) pattern build. The process is currently limited by the maximum feedrate of this CNC router (a maximum of ~350ipm (9m/min)) although the pattern material could be machined faster. Figure 13b presents a closer image of interior of the finished pattern; while Figure 13c shows the resulting sand mold pulled from this pattern cavity.

Figure 13b presents a closer image of interior of the finished pattern; while Figure 13c shows the resulting sand mold pulled from this pattern cavity. This pattern was used to successfully cast a large steel prototype component.~15in (380mm)measuring(a)(b)Figure 13 – Example Pattern and Mold; a) Pattern in process machining through a layer and exposing pattern cavity below, b) finished pattern with complex geometry deep in cavity, and c) chemically bonded sand mold pulled from patternConclusion This paper presented a critical enabling technique in the rapid manufacturing of patterns using an additive/subtractive approach. The research addresses a need for effective layout of layers in this process, as it has been found that layer placement has a significant effect on surface quality of the patterns and more important can avoid catastrophic failure during the machining processes.

The algorithm presented deals effectively with the set of feature conditions that must be addressed. Feature creation is not an inherent problem in this system; since the geometry of sand casting patterns has relatively well-known characteristics. The problem arises in the additive/subtractive nature of the process, as this creates temporary geometric problems such as thin webs of material, potential fracture conditions, etc. However, it is also the layer based nature of the process that enables the rapid prototyping of these patterns, since the process planning is greatly simplified; being able to machine each layer with small tools capable of creating small features and no collision conditions.Acknowledgements Research was sponsored by the U.S. Army Benet Laboratories and was accomplished under Cooperative Agreement Number W15QKN-06-R-0501.

Army Benet Laboratories and was accomplished under Cooperative Agreement Number W15QKN-06-R-0501. The views and conclusions contained in this document are those of the authors and should not be interpreted as representing the official policies, either expressed or implied, of U.S. Army Benet Laboratories or the U.S. Government. The U.S. Government is authorized to reproduce and distribute reprints for Government purposes notwithstanding any copyright notation heron.Reference Binnard, M. and Cutkosky, M. (1998), “Building block design for layered shape manufacturing”, Proceedings-1998ASME Design Engineering Technical Conference , pp. 1-9.Broek, J.J., Horváth, I., Smit, B., Lennings, A.F., Rusák, Z., Vergeest, J.S.M. (2002), “Free-form thick layer objectmanufacturing technology for large-sized physical models”, Autom. in Construction, Vol. 11 No. 3, pp. 335-347.Chang, Y.C., Pinilla, J.M., Kao, J.H., Dong, J., Ramaswami, K., Prinz, F.B.

drocolloid Printing: A Novel Platform for Customized Food Production Daniel L. Cohen1, Jeffrey I. Lipton1, Meredith Cutler3, Deborah Coulter3, Anthony Vesco3, Hod Lipson1,2 1Department of Mechanical and Aerospace Engineering, Cornell University, Ithaca, NY 2Computing and Information Science, Cornell University, Ithaca, NY 3School of Hotel Administration, Cornell University, Ithaca, NYABSTRACTSolid Freeform Fabrication (SFF) of food has the potential to drastically impact both culinary professionals and laypeople; the technology will fundamentally change the ways we produce and experience food. Several imposing barriers to food-SFF have been overcome by recent open- source printing projects. Now, materials issues present the greatest challenge. While the culinary field of molecular gastronomy can solve many of these challenges, careful attention must be given to contain materials-set bloat.

While the culinary field of molecular gastronomy can solve many of these challenges, careful attention must be given to contain materials-set bloat. Using a novel combination of hydrocolloids (xanthium gum and gelatin) and flavor agents, texture and flavor can be independently tuned to produce printing materials that simulate a broad range of foods, with only a minimal number of materials. In addition to extensively exploring future applications of food-SFF, we also present a rigorous proof-of-concept investigation of hydrocolloids for food-SFF. A two-dimensional mouthfeel rating system was created (stiffness vs. granularity) and various hydrocolloid mixtures were characterized via an expert panel of taste testers.INTRODUCTIONFew things are as natively intertwined with humanity as food, which is essential to biological and social life (Counihan, 1999). Not only does food support life and underpin social relations, but it also accounts for a substantial part of our economy.

Not only does food support life and underpin social relations, but it also accounts for a substantial part of our economy. As of 2008, Americans spent $1.02 trillion annually on food, i.e., 9.6% of the nation’s combined disposable personal incomes (USDA Economic Research Service, 2009). Solid Freeform Fabrication (SFF) has the potential to leverage its core strengths (e.g., geometric complexity, automated fabrication) and make its mark on the culinary realm by transforming the way we produce and experience food. We foresee SFF extending new capabilities both to culinary professionals and to laymen, as well as having global social implications by directly extending Web 2.0 phenomena to food.Technological innovations are necessary, however, before these visions can be realized. In addition to lowering barriers to SFF, such as cost of the machine, materials must be developed to feasibly enable a wide range of foods to be produced on SFF platforms.

In addition to lowering barriers to SFF, such as cost of the machine, materials must be developed to feasibly enable a wide range of foods to be produced on SFF platforms. While many raw foods are natively printable, other foods can only be printed if novel materials are developed. Advancements in the field of molecular gastronomy will likely influence the development of these novel food-SFF materials, especially with respect to the use of hydrocolloids. Food-SFF material platforms must be developed in such a way, however, that enables the printing of a very wide range of foods without bloating the size of the required materials-set.

Food-SFF material platforms must be developed in such a way, however, that enables the printing of a very wide range of foods without bloating the size of the required materials-set. A potential solution to the issue of materials-set bloat is the use of novel combinations of hydrocolloids and flavor additives; this platform delivers many degrees of freedom in texture and flavor, while comprising a minimal number of required materials.In this paper, we conceptually explore the broader future implications of food-SFF, as well as present rigorous scientific experiments that tackle materials issues that are central to its feasibility.BACKGROUND: PRIOR WORK AND LOWERING OF KEY BARRIERSThere are several barriers that have been standing in the way of food-SFF. One of the most imposing barriers is that most SFF machines have very limited material-sets.

One of the most imposing barriers is that most SFF machines have very limited material-sets. For example, machines typically only support a handful of materials and nearly none incorporate food-safe printing materials into their repertoire. Moreover, most machines exclusively use proprietary materials and tools, and therefore leave no room for experimentation with food-related printing materials. Thus, it remains that a materials-flexible, open-source printing platform is critical for enabling the growth of the food-SFF paradigm.There are indeed several open-source printing projects that lend the necessary freedom for innovators to experiment with novel, non-proprietary printing materials. One such project is RepRap; however, this open-source printing platform does not offer the deposition tools required to print edible food (Biever, 2005). CandyFab, is another open-source printing project. This platform uses a bed of sugar and a sintering tool to build 3D prototypes (Irwin, 2007).

CandyFab, is another open-source printing project. This platform uses a bed of sugar and a sintering tool to build 3D prototypes (Irwin, 2007). Even so, the primary use is not for producing edible food. Rather than being used for its potentially edible nature, sugar is mainly used because it is a cheap, available and safe printing material. Only one open-source printing platform offers a flexible enough materials-set to enable printing of a wide range of edible products, that is, the Fab@Home. The Fab@Home printing platform is not only open-source, but its syringe-based deposition tool allows for the printing of any material that can be loaded into a syringe (Malone & Lipson, 2007). This non-propriety, flexible-printing- material system overcomes some of the most imposing barriers to food-SFF.Another barrier has traditionally been printing system cost.

This non-propriety, flexible-printing- material system overcomes some of the most imposing barriers to food-SFF.Another barrier has traditionally been printing system cost. While large industrial enterprises may be able to afford $100,000+ SFF systems, many culinary community members and especially individual homeowners cannot get into the food-SFF game with such high equipment costs. In the last decade, we have seen SFF machine costs drop nearly an order of magnitude (with several open-source printers running under $1,500). It is only now, with the reduced SFF cost landscape, that food-SFF is feasible from a financial perspective.To this point, the volume of work published on food printing has been rather limited. Daniel Periard et al investigated SFF of traditional food items that were printable in their raw form (e.g., chocolate, cake frosting).

Daniel Periard et al investigated SFF of traditional food items that were printable in their raw form (e.g., chocolate, cake frosting). This food printing work was limited to those edible printing materials that were natively extrudable through a syringe and inherently held their shape under gravity (Periard, Schaal, Schaal, Malone, & Lipson, 2007).

As a result of relying upon unmodified foods, this work did not fully explore the potential of food printing that could be realized with the development of novel material-sets.Now that the proprietary restrictions, materials-set inflexibilities and cost barriers have been lifted, the stage is set to explore food-SFF’s full potential by inventing new food-related printing applications and by developing the necessary novel materials.POTENTIAL IMPACT OF FOOD-SFFImpact on Culinary Professionals: OverviewFood-SFF would benefit the professional culinary domain primarily in two respects: by lending new artistic capabilities to the fine dining domain, and also by extending mass- customization capabilities to the industrial culinary sector.Impact on Culinary Professionals: Fine DiningFine dining chefs are continually developing new, innovative techniques and seeking the enabling technologies that will help them push the boundaries of culinary art.

In particular, chefs pioneering the cooking style known as “molecular gastronomy” (i.e., the scientific study of physical and chemical processes related to cooking) are at the forefront of this expanding frontier. They innovate by harnessing non-traditional ingredients, such as hydrocolloids, and by employing new tools pulled straight from the scientific community; the result is “culinary magic” including flavored gelatin spheres with liquid centers, sauce foams, hot liquid deserts with flash frozen shells, syringe-extrudable meats, and much more (Allen, 2008). One of the most notable molecular gastronomists is Chef Ferran Adrià, whose restaurant in Spain, elBulli, is considered to be among the finest in the world; the restaurant receives about 2 million applications for a chance at grabbing one of only 8,000 annual slots (Callaway, 2009).

Other notable molecular gastronomists include David Arnold and Nils Noren, of the French Culinary Institute in New York City (Allen, 2008).These pioneers are driving a new trend, in which a growing number of chefs are recognizing the important role that science and technology play in the culinary arts, a notion highlighted by Adrià’s closing of elBulli for 6 months each year to purely focus on R&D (Callaway, 2009). During these closings, Adrià searches the world for the newest enabling technologies that will allow him to push the limits of the culinary arts even further.SFF promises to be the next important enabling technology in the fine dining realm. SFF delivers new possibilities by lending this faction of culinary artists one of SFF’s core capabilities: fabrication of multi-material objects with high geometric complexity.

SFF delivers new possibilities by lending this faction of culinary artists one of SFF’s core capabilities: fabrication of multi-material objects with high geometric complexity. SFF’s culinary potential has already been recognized by one of the most prominent molecular gastronomists; in a 2008 interview with Popular Science, Chef David Arnold mentioned that he “particularly dreams of getting a deal on a 3D rapid prototyping machine” (Allen, 2008). As the barriers fall (e.g., SFF machine prices have reduced nearly an order of magnitude in the last decade) and non-traditional ingredients gain credibility in the fine dining world (e.g., hydrocolloids), the question is not whether SFF will play an important role in the future of food, but rather, in what ways will it do so.Examples of potential future applications include cakes with complex, embedded 3D letters, such that upon slicing the cake, a message is revealed. Or, even a prime rib with a hidden message.

Or, even a prime rib with a hidden message. Perhaps an on-demand, customizable menu in which the dish is prepared in any 3D shape that the diner desires: the diner can co-create with the culinary artist in real-time.Impact on Culinary Professionals: Industrial ProductionThe second way in which SFF could benefit the professional culinary community is by enabling mass-customization in the industrial culinary sector. Today, industrial food producersrely heavily on high-throughput processes such as molding, extrusion and die-cutting (Smith, 2004). These processes, however, are not amenable to mass-customization (i.e., the use of flexible manufacturing techniques to produce custom output in a low-unit-cost fashion). Molding, extrusion and die-cutting each require substantial custom-tooling, and consequently, producing custom output for low-quantity runs is simply unfeasible (Pine, 1993).

Molding, extrusion and die-cutting each require substantial custom-tooling, and consequently, producing custom output for low-quantity runs is simply unfeasible (Pine, 1993). This is precisely where SFF’s inherent strengths can be leveraged: producing food with custom, complex geometries while maintaining cost-effectiveness. The cost-effectiveness is enabled by the fact that SFF does not require custom-tooling or extensive manual labor.One potential future application is custom production of edible giveaways, for example, as marketing collateral for small corporate events. Currently, the cost of custom tooling prohibits low-quantity custom production runs, but with a flexible culinary production platform like SFF, such production runs would be feasible.Impact on Non-Professionals: OverviewCulinary professionals are more primed to adopt SFF than are homeowners, however, the implications for laypeople are even more profound.

The effect on laypeople is essentially twofold: increasing productivity and injecting knowledge.Impact on Non-Professionals: ProductivitySince the late 19th century, a number of machines have been introduced to the home that perform routine tasks more efficiently and ultimately offload work from individuals. Several examples include washing machines, dryers, dishwashers and vacuum cleaners. Perhaps the most profound domestic technology was the sewing machine. Although sewing machines were developed for factory use in the 1850’s, it was not until the late 1870’s that the devices entered homes ubiquitously. Prior to the sewing machine, an average middle-class woman would spend several days per month making and mending her family’s clothing (Museum of American Heritage, 2005). More specifically, it would take 14 hours to make a man’s dress shirt and 10 hours to make a simple dress. With a sewing machine, dress shirts could be made in 1.25 hours and dresses in 1 hour.

More specifically, it would take 14 hours to make a man’s dress shirt and 10 hours to make a simple dress. With a sewing machine, dress shirts could be made in 1.25 hours and dresses in 1 hour. Women of the 1870’s and 1880’s used this saved time to branch out: taking in sewing work for extra money and becoming sales representatives for sewing machines. The reduction in housework ultimately transformed women’s roles as household managers and contributed to women seeking employment outside of the home (Museum of American Heritage, 2005). Domestic technologies have had a tremendous effect on society over the last 150 years by offloading housework, and in turn, creating capacity for people to extend themselves beyond the home in new ways.

SFF has the potential to join this portfolio of important domestic technologies by end-to-end offloading of food preparation.Currently, the average American spends more than 30 minutes per day preparing food, according to USDA economists (Mancino & Newman, 2006). This amount changes drastically depending on a number of factors, including: marital status, working status, gender and income level. Single, working people spend 15 – 35 minutes per day cooking, while married, non- working women spend an average of 70 – 85 minutes per day (Mancino & Newman, 2006). Clearly there is potential to offload culinary housework for people of all genders, marital statuses, working statuses and income levels.

Clearly there is potential to offload culinary housework for people of all genders, marital statuses, working statuses and income levels. If food-SFF were brought to the “set-and-forget” state, requiring minimal human labor, the average person could possibly realize time savings of 150+ hours per year (3.8 workweeks per year), with certain, large groups (e.g., married, non-working women) saving in excess of 500 hours per year (12.5 workweeks per year). Of course, operating the SFF system will still take some finite amount of time and, moreover, families will certainly choose to spend some of the mealtime either with a traditionally cooked meal or out-of- home dining.

Regardless of how much time savings SFF could afford, and even whether it has potential adverse social effects, it is clear that food-SFF is something that warrants further contemplation and investigation.Impact on Non-Professionals: Injecting KnowledgeThe second way that food-SFF could impact laypeople is by abstracting culinary knowledge and injecting it directly into the home. The idea of abstracting knowledge is nothing new. When it comes to playing a popular song, amateur musicians do not have to learn how to play it from scratch. Rather, they obtain sheet music that prescribes the actions (e.g., valve, key or fret manipulations) necessary to reproduce the song. The composer’s artistic skill and knowledge have been abstracted and captured in the sheet music, which somebody lacking the artistic skill and knowledge can use to reproduce the original work.

The composer’s artistic skill and knowledge have been abstracted and captured in the sheet music, which somebody lacking the artistic skill and knowledge can use to reproduce the original work. Of course, the end-user (i.e., the musician) still needs a non-trivial skill set to properly interpret and execute the prescription (i.e., the sheet music).Abstracting knowledge for the purpose of having less skilled practitioners reproduce the original work is also found in the realm of culinary arts. When chefs create new dishes and then write recipes, they are effectively abstracting their knowledge and distilling it into a prescription for others to reproduce their work. Nevertheless, just like the skills a musician needs to effectively play a song from sheet music, a recipe follower still needs non-trivial skills to execute a recipe.It is not only in the abstraction of knowledge, but also in the execution of the prescription that SFF could have tremendous impact.

Just as MIDI software can offload musical skill by taking in digital sheet music and directly creating sound, the SFF system could directly inject the skills necessary to follow a recipe end-to-end. Laypeople don’t have to know the first thing about musical notation, valve/key/fret fingering, or tonal theory to be able to utilize a stereo system to deliver a distilled version of a live musical performance directly into their home. Likewise, a layperson would not necessarily need to possess even basic culinary skills to employ an SFF system to create geometrically complex, multi-material food items.Culinary knowledge and artistic skill of world renowned chefs can be abstracted to a 3D fabrication file and then used by laypeople to reproduce famous chefs’ work in the home.

Also, expert knowledge of the world’s leading nutritionists can be abstracted and encoded in 3D fabrication files to help laypeople eat more healthily, without necessarily having to learn healthy cooking techniques or even understand nutritional principles such as caloric intake and protein balance. SFF systems could even go one step further, and deliver customized solutions (SFF’s core strength) to each user that incorporate the individualized nature of nutritional needs. For example, a layperson may soon be able to upload a report of their daily activity from a pedometer and digital food log, and the SFF system could use expert knowledge to print them a meal that fulfills their particular nutritional needs for the day.

While experts can currently offer advice on how to balance a nutritional program, their influence falls short of delivering the end- to-end solution that only SFF system can provide: from personalized design through fabrication.Food-SFF and Web 2.0Whether professionals use SFF for cutting-edge capability or laypeople use it to borrow more basic skills, the abstraction and subsequent direct-execution of culinary knowledge has profound social effects. Once culinary knowledge can be abstracted, it can also be easily shared. It wasn’t until music and photographs were represented in digital abstracts (i.e., MP3s and JPEGs) that sharing became easy enough for entire web communities to develop around them.Some online Web 2.0 social networks have grown to more than 200,000,000 community members (Powell, 2009). In particular, networks established around sharing and communally experiencing photographs (e.g., Flickr) and music (e.g., Last.fm) boast tens of millions of users, each.

In particular, networks established around sharing and communally experiencing photographs (e.g., Flickr) and music (e.g., Last.fm) boast tens of millions of users, each. Now that SFF will potentially enable the feasible sharing of food (via trading 3D fabrication files), we can start thinking of what the future could look like. After all, few things are more essential to social relationships than food. Through food, people share tastes, ideas, values and generosity (Counihan, 1999). Imagine amateurs and professionals alike, sharing their latest ideas and helping others solve culinary challenges. Not only would new types of social bonds form across the world, but people will harness principles of “wikinomics” and begin to “mass-collaborate” seamlessly from all corners of the globe (Tapscott & Williams, 2006). A truly global cuisine could emerge.

A truly global cuisine could emerge. Through mass-collaboration, people could even attempt to solve major challenges such as creating fundamentally new types of healthful foods.Furthermore, we could see democratization of innovation and a major revision of the current business models of the culinary profession. No longer would only those with access to distribution channels be privileged with sharing their food with the world, but rather, any amateur or professional chef could see their work gain prominence as long as their ideas have merit. A more democratic business landscape will emerge, and the lines between amateur and professional will be blurred.MATERIALS CHALLENGESIn order to unlock food-SFF’s potential, we need to think about the specific technical challenges standing in the way.

As mentioned earlier, SFF systems’ materials inflexibility, prohibitive cost and proprietary restrictions were barriers that have largely been overcome by recent open-source printing efforts, such as Fab@Home. Now, the printing materials themselves are the bottleneck. Some foods, such as cake frosting, processed cheese, hummus and chocolate are natively printable; they are extrudable through a syringe tip and hold their shape under gravity (Periard, Schaal, Schaal, Malone, & Lipson, 2007). Other foods, such as fruits, vegetables and meats are not natively printable. In order to be able to print these important food- types, we would have to undertake substantial reformulation efforts. These types of challenges have already been tackled by molecular gastronomists. In this avant-garde culinary field, it is becoming typical to make solids (e.g., meats) extrudable by adding hydrocolloids (Allen, 2008).

In this avant-garde culinary field, it is becoming typical to make solids (e.g., meats) extrudable by adding hydrocolloids (Allen, 2008). With the appropriate molecular gastronomic tricks, we can realize both printable solid foods and printable semi-solid liquids; the possibilities are nearly limitless.However, as we attempt to target foods one-by-one in an effort to make them printable, the number of custom, one-off material combinations greatly expands. This leads to the challenge of containing the bloat of the materials-set. The question becomes: how do we enable the printing of a wide range of foods with only a limited, fundamental set of materials?FLEXIBLE MATERIALS PLATFORM: A MOLECULAR GASTRONOMIC APPROACHOverviewOne potential solution to containing materials-set bloat is to use a small group of ingredients to create a platform with many degrees of freedom in terms of texture and flavor.

By fine tuning hydrocolloids’ concentrations, and focusing on the ratios between combinations of different hydrocolloids, a very wide range of textures (i.e., mouthfeels) can be achieved. Furthermore, given the generally neutral flavors of hydrocolloids, the flavor can be independently tuned by using concentrated flavoring additives.

Furthermore, given the generally neutral flavors of hydrocolloids, the flavor can be independently tuned by using concentrated flavoring additives. By independently controlling these two parameters, texture and flavor, we suggest that a wide range of food experiences can be simulated.ChocolateFried fishCarrotsBiscottimr iF t s o MMushroomAppleCooked chickenSaltineBananaCooked spaghettiFresh mozzarellaTomatoHard-boiled egg yolkCreamy peanut butterMarshmallow fluffMeringueGround beefJell-OGravyMashed potatoCake icingRefried beansSelf- supporting gelWhite breadPolentat s e k a e WNon-self- supporting gelMilkPuddingIce creamCake batterRaspberry coulisApple sauceSorbetRisottoCoffee grindsSmoothestMost GranularTable 1 Mouthfeel matrix with common foods placed as reference itemsThe two hydrocolloids focused on in this paper are xanthan gum and gelatin.

Various combinations of the two were tested for mouthfeel. In order to rigorously describe the mouthfeel of each resultant material, we devised a rating system in which the material was rated along two orthogonal axes: 1) weak to firm, and 2) smooth to granular. This “mouthfeel matrix” was developed in cooperation with expert chefs from the Cornell University School of Hotel Administration (Table 1).There are 7 discrete buckets on the smooth-to-granular axis and 9 buckets on the weak-to- firm axis. Common foods were placed within this framework for the taste testers’ reference, and the testers were asked to place a particular hydrocolloid concoction within the matrix relative to these references.Methods: Materials PreparationXanthan gum and gelatin were prepared in water at various concentrations, ranging from 0.5% gelatin in water (by weight) to 4% gelatin in water, and from 2% xanthan in water to 16% xanthan in water.

After preparation, the hydrocolloid concoctions were loaded into 10 mL syringes.Figure 1 Various hydrocolloid formulations loaded into 10 mL Luer-Lok syringesTo test the flexibility afforded by flavoring additives, various food grade flavor concentrateswere infused with the hydrocolloids, including raspberry, strawberry, banana and chocolate.Methods: PrintingAfter the materials were loaded into 10 mL Luer-Lok syringes, the syringes were loaded into the displacement deposition tool of the Fab@Home printer (Figure 2). Using the custom, open- source Fab@Home control software, 3 centimeter cubes were printed.Figure 2 (left) Fab@Home printing platform, (right) Fab@Home control softwareWe printed each sample to properly account for any effect the extrusion process had on the texture.

There is gel-ripping upon extrusion that creates a specific microstructure and texture.Methods: Materials EvaluationVarious combinations of xanthan gum and gelatin were prepared and presented to a panel of seven taste testers, including two professional culinary instructors.

Each rated the hydrocolloids within the “mouthfeel matrix” by comparing the texture to the matrix’s reference foods.Results: Platform Flexibility Described Through Mouthfeel MatrixThe mouthfeels for xanthan gum and gelatin hydrocolloid concoctions were individuallytested and pinned on the mouthfeel matrix.mr i F t s o M4% gelatin CLOSE TO CHOCOLATE/ MUSHROOM16% xanthan COOKED SPAGHETTI1% gelatin 8% xanthan CLOSE TO TOMATO2% gelatin JELL-O2% gelatin 8% xanthan CLOSE TO CAKE ICING/ MERINGUE0.5% gelatin 8% xanthan CLOSE TO CAKE ICING/ MERINGUE4% xanthan MASHED POTATO1% gelatin SELF- SUPPORTING LOOSE GEL 2% xanthan NON-SELF- SUPPORTING LOOSE GEL0.5% gelatin 4% xanthan APPLE SAUCE1% gelatin 4% xanthan RISOTTOt s e k a e W0.5% gelatin MILKSmoothestMost GranularTable 2 Mouthfeel matrix of hydrocolloid mixture showing the formulations in the appropriate locations relative to common foods (see Table 1 for more detail) with the closest common foods are listed below the hydrocolloid concentrations in boldMouthfeels were achieved that simulated a broad range of common foods.

These simulated mouthfeels ranged from milk to mushroom to tomato to risotto. The hydrocolloids spanned a range from liquids to solid vegetables. The mouthfeels followed a rather identifiable pattern. Pure xanthan and gelatin tracks directly up the weak to firm axis, but does not shift in granularity. However, as xanthan and gelatin are combined, the resultant hydrocolloids begin to possess some amount of granularity. Generally, the higher the concentration of xanthan and gelatin, the firmer and more granular the gels become.The printability of the hydrocolloids reached a limit as the stiffness and granularity of theresultant materials prevented reliable extrusion of the materials through a 3 mm syringe orifice.DISCUSSION AND CONCLUSIONSUsing only two ingredients, xanthan and gelatin, a very broad range of mouthfeels can be simulated.

This type of approach addresses the issue of materials-set bloat that would be faced as end-users attempt to feasibly implement SFF for culinary applications with a small, practical materials-set.While the work herein serves as a proof-of-concept, further materials development is required to progress food-SFF. Not only does the mouthfeel range need to be expanded, but the flavors also need to be refined by tuning the flavor additives. To truly demonstrate the capability of hydrocolloid-based printing, ideally a double-blind tasting should be performed and it would be seen whether the taster can distinguish between the natural food and the hydrocolloid version.

It should be noted, however, that even if subtle differences are perceptible, it is not necessary in all cases to perfectly reproduce the original food; there is still great value in simulating the original food.Regardless of whether a hydrocolloid approach is taken to food-SFF, or some other molecular gastronomic platform is employed, the potential future applications of food-SFF remain the same. From culinary professionals to laypeople, individuals from all walks of life will be drastically affected by food-SFF. Artistic boundaries will be pushed in fine dining and industrial producers will explore mass-customization. Laypeople will have housework time reduced and benefit from direct culinary skill injections.

Laypeople will have housework time reduced and benefit from direct culinary skill injections. Web 2.0 will tackle the next great frontier as people from all over the world experience food in new ways, while forming social bonds and mass-collaborating.Now that major barriers have been broken, such as high printer cost and proprietary restrictions, the stage is finally set for tremendous growth of food-SFF. Few things are more central to humanity than food, and therefore is should come as no surprise when food-SFF gains prominence as one of the 21st century’s important domestic technologies.ACKNOWLEDGMENTSThe authors would like to thank the Cornell University School of Hotel Administration,and in particular Chef White for his participation in our panel of expert taste testers.REFERENCESAllen, T. (2008, July 28). Doctor Delicious. Popular Science .Biever, C. (2005, March 18). 3D Printer to Churn Out Copies of Itself. New Scientist Magazine .Callaway, E. (2009, March 20).

DROGELS IN STEREOLITHOGRAPHYKarina Arcaute*, Luis Ochoa*, Brenda Mann**, Ryan Wicker* *University of Texas at El Paso, W.M. Keck Border Biomedical Manufacturing and Engineering Laboratory, Department of Mechanical and Industrial Engineering, El Paso, Texas 79968-0521 **Sentrx Surgical, Inc., Salt Lake City, UTAbstract Reviewed, accepted August 23, 2005The use of stereolithography (SL) for fabricating complex three-dimensional (3D) tissue engineered scaffolds of aqueous poly(ethylene glycol) (PEG) hydrogel solutions is described. The primary polymer used in the study was PEG-dimethacrylate (PEG-dma) with an average molecular weight (MW) of 1000 in distilled water with the photoinitiator Irgacure 2959 (I-2959).

The primary polymer used in the study was PEG-dimethacrylate (PEG-dma) with an average molecular weight (MW) of 1000 in distilled water with the photoinitiator Irgacure 2959 (I-2959). Successful layered manufacturing (LM) with embedded channel architecture required investigation of the photopolymerization characteristics of the PEG solution (measured as hydrogel thickness or cure depth) as a function of photoinitiator concentration and laser energy dosage for a specific photoinitiator type and polymer concentration in solution. Hydrogel thickness was a strong function of PI concentration and energy dosage. Curves of hydrogel thickness were utilized to successfully plan, perform, and demonstrate layered manufacturing of highly complex hydrogel scaffold structures, including structures with internal channels of various orientations.

Successful fabrication of 3D, multi-layered bioactive PEG scaffolds containing cells was accomplished using a slightly modified commercial SL system (with 325 nm wavelength laser) and procedure. Human dermal fibroblast (HDF) cells were encapsulated in PEG hydrogels using small concentrations (~ 5 mg/ml) of acryloyl-PEG-RGDS (MW 3400) added to the photopolymerizable PEG solution to promote cell attachment. HDF cells were combined with the PEG solution, photocrosslinked using SL, and successfully shown to survive the fabrication process. The combined use of SL and photocrosslinkable biomaterials such as PEG makes it possible to fabricate complex 3D scaffolds that provide site-specific and tailored mechanical properties (i.e., multiple polymer materials) with a polymer matrix that allows transport of nutrients and waste at the macroscale and facilitates cellular processes at the microscale through precisely placed bioactive agents.1.

IntroductionWhen stereolithography (SL), a rapid prototyping (RP) or layered manufacturing (LM)process that could automatically build detailed three-dimensional (3D) parts, became commercially available in the mid-1980s, the manufacturing industry was profoundly impacted by significantly improving the way in which prototypes were fabricated (Wohlers, 2004). Commercial line-scan SL solidifies thin layers of photo-reactive liquid polymer using an ultraviolet (UV) laser that selectively crosslinks the liquid photopolymer by rastering across the surface of the polymer contained in a vat. 3D shapes are built one layer at a time by attaching the desired part to a build platform that begins at the surface of the liquid photopolymer and traverses deeper into the liquid after each successive layer.

Since the introduction of SL, a number of new RP technologies have been developed and commercialized with widely varying applications in engineering design, manufacturing, medicine, and more.During the last few years, RP technologies have been employed quite extensively in the fieldof tissue engineering (TE) for manufacturing 3D scaffolds that provide the mechanical support and structure for cells to grow and regenerate damaged tissue. Researchers have investigated a wide variety of strategies for engineering implantable scaffolds for tissue regeneration. The434manufacturing processes used for engineering implantable scaffolds depend on the material used to fabricate the scaffold and the intended application. Porous polymer scaffolds that act as substrates for cell attachment are produced using techniques such as fiber bonding, membrane lamination, and solvent casting-particulate leaching (Lu and Mikos, 1996; Ma, 2004).

However, complex scaffolds that mimic tissue structures have been difficult to produce, or the fabrication method has been difficult to scale up to a manufacturing level. Also, using traditional manufacturing processes has not allowed the fabrication of scaffolds with predetermined internal structures that provide the cells with an optimum local microenvironment. Precise control of the scaffold microenvironment allows the incorporation of the necessary biochemical stimulus for the cells to proliferate, differentiate, migrate, and undergo apoptosis. RP technologies are increasingly demonstrating the potential for fabricating 3D structures with precise control of the micro- and macro-scale characteristics (Liu and Bhatia, 2004). Some researchers have developed new RP technologies for the fabrication of tissue scaffolds based on the 3D dispensing of aqueous solutions and pastes (Landers et al., 2002; Ang et al., 2002).

Some researchers have developed new RP technologies for the fabrication of tissue scaffolds based on the 3D dispensing of aqueous solutions and pastes (Landers et al., 2002; Ang et al., 2002). Other researchers (Cooke et al., 2002; Dhariwala et al.,2004; Comeau et al. 2005) have explored the capabilities of SL for manufacturing biodegradable polymeric scaffolds. Using SL, Cooke et al. (2002) were able to fabricate acellular TE scaffolds using poly(propylene fumarate) (PPF) with applications in bone regeneration where the scaffold contained 3D macroscale features and interconnecting microscale porous networks. In addition to the PPF, a class of hydrophilic polymer materials, called hydrogels, exists that are widely used in tissue engineering and can be photocrosslinked. Poly(ethylene glycol) (PEG) is a synthetic hydrogel polymer used in tissue regeneration because of its characteristics as being considered non-toxic, non-immunogenic and easily cleared from the body.

Poly(ethylene glycol) (PEG) is a synthetic hydrogel polymer used in tissue regeneration because of its characteristics as being considered non-toxic, non-immunogenic and easily cleared from the body. PEG is water soluble and can be modified with photoreactive and crosslinkable groups like acrylates or methacrylates (Harris and Zalipsky, 1997). Further, PEG can be made bioactive by incorporating adhesion ligands, growth factors, and cytokines to the polymer solution during the photocrosslinking process. Bioactive factors can be attached to the scaffold through functional groups or may simply be trapped within the hydrogel during crosslinking (Mann et al., 2001a and b).

Bioactive factors can be attached to the scaffold through functional groups or may simply be trapped within the hydrogel during crosslinking (Mann et al., 2001a and b). These characteristics make PEG a strong candidate for creating photocrosslinkable hydrogel tissue engineered scaffolds (Arcaute et al., 2005b).Our research group has been developing new multiple material SL fabrication technologies for use in TE applications (Wicker et al., 2004) and investigating the use of PEG in SL for TE applications (Arcaute et al., 2005a; Arcaute et al., 2005b; Pallante et al., 2005). Recent attention in peripheral nerve regeneration has focused on developing multi-material and multi-lumen PEG-based nerve guidance conduits (NGCs) with precisely placed bioactive agents and cells to regenerate peripheral nerves faster and over longer distances.

SL has demonstrated that it has the capability of accurately manufacturing these multi-material and multi-lumen NGCs with the additional advantage of being able to scale up for mass production (Arcaute et al., 2005b). As a result, the present work describes our experiences with using SL for fabricatingcomplex 3D tissue engineered PEG-based scaffolds. The primary polymer used here was PEG- dimethacrylate (PEG-dma) with an average molecular weight (MW) of 1000 in distilled water with the photoinitiator Irgacure 2959 (I-2959). The photopolymerization characteristics of the PEG solution were investigated where hydrogel thickness or cure depth was measured as a function of photoinitiator concentration and laser energy dosage. Curves of hydrogel thickness were utilized to successfully plan, perform, and demonstrate layered manufacturing of highly complex hydrogel scaffold structures, including structures with internal channels of various orientations.

Successful fabrication of 3D, multi-layered bioactive PEG scaffolds containing435cells was accomplished using a slightly modified commercial SL system and procedure. Human dermal fibroblast (HDF) cells were encapsulated in the PEG matrix during SL and tested for cell viability. In the cell encapsulation experiments, small concentrations (~ 5 mg/ml) of acryloyl- PEG-RGDS were incorporated in the photopolymerizable PEG solution to fabricate bioactive scaffolds that promote cell attachment. The following sections describe the methods and results of these investigations.2. Experimental Methods2.1. Cure Depth ExperimentsInvestigation of the cure depth characteristics of a hydrogel material as a function of differentfactors was carried out previously (Arcaute et al., 2005a).

Cure Depth ExperimentsInvestigation of the cure depth characteristics of a hydrogel material as a function of differentfactors was carried out previously (Arcaute et al., 2005a). In this earlier study, cure depth, hydrogel thickness, or simply gel thickness was found to be a strong function of photoinitiator (PI) type and concentration, energy dosage, and polymer concentration in the solution. For the two PIs considered in the previous study, Irgacure 2959 (I-2959) appeared to be a superior candidate to use in the fabrication of scaffolds in a layer-by-layer fashion. Additional cure depth experiments were carried out and are described in the following sections to better understand the photocrosslinking process of a hydrogel material.2.1.1. MaterialsThe polymer used was PEG-dma with average MW of 1000 (Cat. No. 15178, Polysciences, Inc., Warrington, PA). PEG was chosen because it is a synthetic material with important use in tissue regeneration.

No. 15178, Polysciences, Inc., Warrington, PA). PEG was chosen because it is a synthetic material with important use in tissue regeneration. The PI used was I-2959 (2-hydroxy-1-[4-(2-hydroxyethoxy)phenyl]-2- methyl-1-propanone) (PN 1559163, Ciba Speciality Chemicals Corp., Tarrytown, NY). Cytocompatibiblity studies (Bryant et al., 2000; Leach et al., 2003) have shown that I-2959 is an advantageous radical-forming photoinitiator in cell encapsulation applications since I-2959 does not have a significant detrimental effect on cell survival. Also, I-2959 allows the photopolymerization of PEG-based solutions as thin polymer networks that permit the fabrication of 3D PEG scaffolds using SL (Arcaute et al., 2005a). The polymer was dissolved in distilled water to prepare a 30% (w/v) solution. Aliquot portions were separated from this solution, and different concentrations of I-2959 were added.

The polymer was dissolved in distilled water to prepare a 30% (w/v) solution. Aliquot portions were separated from this solution, and different concentrations of I-2959 were added. The PEG-dma solution with PI is the photoreactive solution used in the cure depth experiments.2.1.2. ApparatusA 3D Systems Model 250/50 SL machine equipped with a He-Cd laser (325 nm) was used inthis study. The original setup of the SL machine was modified slightly to carry out the gel thickness experiments. The modifications consisted of removing the original vat of material from the machine and setting the elevator platform at a fixed height. A flat-top cylindrical container was placed on the center of the platform as depicted in Figure 1. At the test height of the elevator platform, the laser beam was circular with a diameter of ~250 µm.4362.1.3. ProcedureThe PEG-dma solution with PI was pipetted inside the container and filled up to the rim.

ProcedureThe PEG-dma solution with PI was pipetted inside the container and filled up to the rim. Aglass slide was placed on top of the container and in contact with the PEG-dma solution. The glass slide acted as a substrate for gel attachment and facilitated measurement of gel thickness. Samples were cured by writing a vector pattern through the glass slide and into the PEG-dma container at different energy dosages. The vector pattern consisted of a series of nineteen parallel lines with each line ~250 µm wide, or one laser beam width, 7.62 mm long and spaced ~355 µm apart. This vector pattern formed a ~6.64 mm by 7.62 mm rectangle (in the x and y dimensions) from which the sample cure depths were measured (in z).

This vector pattern formed a ~6.64 mm by 7.62 mm rectangle (in the x and y dimensions) from which the sample cure depths were measured (in z). The laser beam was commanded to draw the vector pattern, and the energy dosage was varied by changing the SL machine parameters that control the scan speed of the laser (Arcaute et al., 2005a).After polymerization, the glass slide was lifted off the container with the polymerized gel attached to the slide. The cured gels were flushed with distilled water to remove unreacted PEG- dma solution, and the thickness (in z) of the gels was then measured with a caliper. Five sample gels were cured for each of the PEG-dma solutions, and the results presented in the next section represent the average thickness measurements for the samples. It was determined by using a laser power meter that the glass slide filtered the laser power by ~18%.

It was determined by using a laser power meter that the glass slide filtered the laser power by ~18%. The measured power of the He-Cd laser (rated at 40 mW) at the fixed platform height was 29.5 to 30.6 mW and 35.8 to 37.1 mW with and without the glass slide, respectively. The measured laser powers with the glass slide along with the speed of the laser beam given by the SL machine were used to determine the laser energy dosages.2.2. Scaffold FabricationThe modified SL machine used in the cure depth experiments was used for the fabrication ofPEG-based scaffolds. A flat-bottom container is used as the receptacle for the photoreactive solution instead of the original vat. In this modified setup, a thin (0.45 mm) glass slide was placed within the flat-bottom container to act as a platform for the scaffolds and facilitate their removal once the scaffolds were built.

In this modified setup, a thin (0.45 mm) glass slide was placed within the flat-bottom container to act as a platform for the scaffolds and facilitate their removal once the scaffolds were built. The elevator platform started at a height at which the laser beam was circular with a diameter of ~250 µm, and the receptacle contained sufficient photoreactive polymer to form a single layer. The elevator was lowered every time moreHe-Cd LaserFocusing OpticsScanning MirrorsGlass SlideMovable PlatformFlat-top Container w/PEG-dma SolutionFigure 1. Schematic of SL showing the setup to perform the gel thickness experiments.437polymer solution was added to the flat-bottom container in order to construct a new layer and maintain the laser beam diameter at ~250 µm. Fabrication of multiple material scaffolds is also feasible using this setup by adding and withdrawing materials as each layer is fabricated.2.3.

Fabrication of multiple material scaffolds is also feasible using this setup by adding and withdrawing materials as each layer is fabricated.2.3. Fabrication of Hydrogel Scaffolds Containing Live CellsLive cells and bioactive agents were combined with PEG-dma photoreactive solution and photocrosslinked using SL. For these experiments, human dermal fibroblasts (HDF) cells were encapsulated in bioactive PEG scaffolds. The tetrapeptide RGDS was conjugated to PEG and added in small concentrations to the photoreactive solution to promote cell adhesion. The following sections describe the experimental methods and procedures carried out to create cell- containing scaffolds.2.3.1.

The following sections describe the experimental methods and procedures carried out to create cell- containing scaffolds.2.3.1. Cell MaintenanceHuman dermal fibroblasts (HDFs) were obtained from Cambrex Bio Science (Walkersville, MD) and maintained on Dulbecco’s Modified Eagle Medium (DMEM) supplemented with 10% Fetal Bovine Serum (FBS; BioWhittaker, Walkersville, MD), 200 mM L-glutamine, 10,000 units/mL of penicillin and 10 mg/mL of streptomycin. Cells were incubated at 37ºC in a 5% CO2 environment. Cells at passage four were used in the experiments after 80-90% confluence was reached.2.3.2. Peptide Conjugation to PEGThe tetrapeptide Arg-Gly-Asp-Ser (RGDS) was conjugated to PEG by reacting the peptide with acryloyl–PEG–N-hydroxysuccinimide (acryloyl–PEG–NHS, 3400 Da; Nektar Therapeutics, Huntsville, AL) in 50 mM sodium bicarbonate buffer (pH 8.5) for 2 h. The mixture was then lyophilized and stored frozen.

The mixture was then lyophilized and stored frozen. The RDGS sequence is a known cell adhesion ligand (Mann et al., 2001; Mann and West, 2002). The resulting acryloyl-PEG-RGDS was used to promote cell attachment and fabricate bioactive scaffolds.2.3.3. Bioactive Scaffold FabricationFirst, the PEG-dma and the photoinitiator I-2959 were dissolved in a 10 mM HEPES (N-(2- Hydroxyethyl)piperazine-N′-(2-ethanesulfonic acid) buffered saline solution (pH 7.4). A small concentration (5-10 mg/mL) of acryloyl-PEG-RGDS was then added to the photoreactive PEG- buffered solution, and the solution was filter sterilized. A cell suspension in complete media was mixed in equivolume amount with the PEG solution to form a solution with the following concentrations: 0.25 x 106 cells/mL, 30% (w/v) PEG-dma, 0.5% I-2959, and 5 mg/mL of acryloyl-PEG-RGDS.

The final PEG-cells solution was then pipetted inside the flat-bottom container, and the SL machine was commanded to photocrosslink a simple pattern (ID = 3 mm, OD = 5mm). The scaffolds were fabricated with three different conditions: (a) single layer (1- mm thick), (b) single layer – pattern drawn twice (1-mm thick), and (c) double layer (2-mm thick). Eight scaffolds were fabricated for each condition in batches of four scaffolds.Once the cell-containing scaffold was fabricated, it was placed in a 96-well plate with mediaand placed in an incubator at 37ºC and 5% CO2. Cell viability of the encapsulated cells was tested with the LIVE probe of the LIVE/DEAD® assay (Molecular Probes, LIVE/DEAD® Viability/Cytotoxicity Kit L-3224, Eugene, OR) at 48 hours. To perform the viability assay the old media was removed, and the gels were rinsed twice with Dulbecco’s phosphate buffered saline solution (D-PBS).

To perform the viability assay the old media was removed, and the gels were rinsed twice with Dulbecco’s phosphate buffered saline solution (D-PBS). After that, the gels were stained with Calcein AM 0.4 µM and observed using a confocal microscope with fluorescence capabilities.4383. Results3.1. Cure Depth CurvesThe experimental measurements of gel thickness as a function of PI concentration and energy dosage are shown in Figures 2 and 3. The gel thickness curves as a function of PI concentration shown in Figure 2 illustrate that the gel thickness begins at zero for zero PI concentration, has a maximum at low PI concentrations, and decreases asymptotically to a non-zero value as PI concentrations increase. The gel thickness starts at zero due to the presence of polymerization inhibitors, including monomethyl ether hydroquinone (MEHQ) and butylated hydroxytoluene (BHT), added by the manufacturer.

The gel thickness starts at zero due to the presence of polymerization inhibitors, including monomethyl ether hydroquinone (MEHQ) and butylated hydroxytoluene (BHT), added by the manufacturer. Measurements in the region between zero and the maximum gel thickness are not presented here as the gel is loosely crosslinked at these PI concentrations and the measurements are highly uncertain (and thus, the maximum gel thickness presented here should not be viewed as an absolute maximum). The level of crosslinking is also related to the PI concentration.

The level of crosslinking is also related to the PI concentration. In general, for a given laser energy, low PI concentrations correspond to thicker, loosely crosslinked gels, while high PI concentrations correspond to highly crosslinked and thinner gels.18181616)m m( s s e n k c i h T1412108)m m(s s e n k c i h T1412108l e G64l e G6422000.25 0.5 0.7511.25 1.5 1.7522.25 2.5 2.7530 0.010.1110Photoinitiator Concentration (% w/v)Photoinitiator Concentration (% w/v)Figure 2. Gel thickness curves versus photoinitiator concentration for a 30% (w/v) PEG-dma solution in distilled water: linear scale (left) and log scale (right). The three different types of markers correspond to three different energy dosages (♦ 3.604 J/cm2 [0.258 IPS], ■ 1.640 J/cm2 [0.567 IPS], ▲ 0.586 J/cm2 [1.585 IPS]). Error bars not shown are contained within the symbol.The energy dosage significantly affects gel thickness.

Error bars not shown are contained within the symbol.The energy dosage significantly affects gel thickness. High energy dosages generallyproduce thicker gels. Figure 3 reports the relationship between energy dosage and gel thickness, for a PEG solution with 0.5% (w/v) I-2959. For PI concentrations smaller than 0.05%, small energy dosages (fast scanning speeds) produce thick and loosely crosslinked gels.3.2. Scaffold Fabrication DemonstrationComplex 3D hydrogel structures were fabricated using the cure depth data by selecting a combination of layer thickness and energy dosage (scan speed and laser energy) that provided adequate resolution for the desired part and produced sufficiently crosslinked gels, respectively. The pictures in Figure 4 show two chess rooks manufactured using SL. The rook on the right in each picture was manufactured using the 3D Systems Viper™ Si2 SL apparatus using DSM Somos® WaterShed™ 11120 resin.

The rook on the right in each picture was manufactured using the 3D Systems Viper™ Si2 SL apparatus using DSM Somos® WaterShed™ 11120 resin. The rook on the left in each picture was built in the modified SL 250/50 apparatus with a 30% (w/v) PEG-dma solution with 3% (w/v) of I-2959. As4391.61.61.41.41.21.2)m m1)m m(1s s e n k c i h Tl e G0.80.6s s e n k c i h Tl e G0.80.60.40.40.20.20005001000150020002500110100100010000Energy Dosage (mJ/cm2)Energy Dosage (mJ/cm2)Figure 3. Gel thickness vs. energy dosage for a 30% (w/v) PEG-dma solution with 0.5% (w/v) of I-2959: linear scale (left) and log scale (right).shown in the picture in the middle of Figure 4, the PEG-based chess rook is elastic and can be easily deformed. The PEG-based chess rook includes all the details of the resin-based rook such as the staircase contained within the walls, the windows, and the bricks in the facade.

The PEG-based chess rook includes all the details of the resin-based rook such as the staircase contained within the walls, the windows, and the bricks in the facade. The PEG- based piece measures approximately 23 mm, and was fabricated in 0.010-in layers. The picture on the right illustrates the shrinkage experienced by PEG when dry.Other complex PEG-based structures with embedded channels can also be fabricated using SL. The pictures in Figure 5 show a block with a single embedded channel that bifurcates into two branches. These two branches then bifurcate again into two channels (each) in a different plane of the gel (each bifurcation is oriented 90o from one another). This bifurcating geometry is representative of the arterial system and a gel such as this one may potentially be used as a tissue construct in angiogenesis studies (along with the appropriate bioactive agents and cells).

The dimensions of the PEG block are 1 by 1 by 1.6 cm, and the diameters of the embedded channels are 1 mm. The scaffold was built with a 30% (w/v) PEG solution containing 0.5% I-2959. Scaffolds of virtually any length and geometry can be fabricated as demonstrated by the scaffold in the bottom of Figure 5, which starts as a single channel and bifurcates twice with an overall length of approximately 5 cm.Figure 4. Complex 3D structures manufactured using SL, the chess rook on the right in each picture was made with DSM Somos® WaterShed™ 11120 resin while the rook on the left is PEG-based. The picture on the left shows a PEG-based structure immediately after it was built. The picture in the center shows how easily a PEG-based structure can be deformed. The picture on the right shows a dry PEG-based structure, illustrating the shrinkage experienced by PEG when dry.440Figure 5.

The picture on the right shows a dry PEG-based structure, illustrating the shrinkage experienced by PEG when dry.440Figure 5. PEG-based scaffolds with embedded channels of different sizes and orientations: isometric view (top left) side view (top right) of a 1 by 1 by 1.6 cm block, isometric view of a 1 by 1 by 5 cm block.The scaffolds shown in Figure 6 illustrate examples of multi-material fabrication. Each different color corresponds to two layers of 1 mm thickness each. The design corresponds to a multi-lumen nerve guidance conduit (NGC) with end caps for suturing to the severed nerve stumps. The dimensions of the end caps are OD = 5 mm, ID = 3 mm, and the multi-lumen section contains twelve lumens, each 500 µm in diameter. Commercially available NGCs are artificial tubes, made out of biocompatible polymers such as silicone and polyglycolic acid (PGA) and their water content is similar to human tissue.

Commercially available NGCs are artificial tubes, made out of biocompatible polymers such as silicone and polyglycolic acid (PGA) and their water content is similar to human tissue. These single-lumen NGCs are designed for permanent implantation (in the case of PGA, the polymer is eventually bio- absorbed) and have proven to be effective in repairing damaged nerves. Occasionally these tubes require removal due to local discomfort caused by inflammation at the severed nerve ends (Lundborg et al., 2004). Research in TE nerve regeneration is now introducing bio-absorbable, bio-degradable materials, growth factors, and cells into the NGCs to promote faster healing over longer lengths. Current research is examining conduit designs with multiple lumens in order to better mimic the natural structure of nerves (Hadlock et al., 2000, Sundback et al., 2003, Lundborg et al., 2004).

Current research is examining conduit designs with multiple lumens in order to better mimic the natural structure of nerves (Hadlock et al., 2000, Sundback et al., 2003, Lundborg et al., 2004). A low-pressure injection molding, thermally induced phase transition technique was recently developed to fabricate multi-lumen biodegradable NGCs (Sundback et al., 2003). These five-lumen conduits are made out of polylactide-co-glycolide (PLGA) and coated with laminin to promote the adhesion of Schwann cells that are seeded into the scaffolds prior to implantation in rats. After six weeks, in vivo regeneration in these conduits was favorably compared with autografts. Even though these conduits were shown to be effective in the regeneration of nerves, the low-pressure injection molding fabrication technique would be difficult to scale-up to a manufacturing level.

Even though these conduits were shown to be effective in the regeneration of nerves, the low-pressure injection molding fabrication technique would be difficult to scale-up to a manufacturing level. The combined use of SL and photocrosslinkable biomaterials (like PEG) makes it possible for the bulk fabrication of complex 3D scaffolds. In the case of NGCs, SL successfully fabricates multi-lumen and multi-material NGCs that along with precisely placed bioactive agents and cells could be used to regenerate peripheral nerves faster and over longer distances. Our group is investing considerable effort in this area of research at the present time (Arcaute et al., 2005b).441Figure 6. PEG-based multi-lumen scaffolds containing different colors of 15-µm fluorescent particles: isometric view (left), side view (center), and top view (right).3.3.

PEG-based multi-lumen scaffolds containing different colors of 15-µm fluorescent particles: isometric view (left), side view (center), and top view (right).3.3. Scaffolds Containing Live Cells and Bioactive FactorsHDF cells encapsulated inside PEG scaffolds fabricated in SL successfully survived the fabrication process. The photocrosslinked samples containing cells were stained with the LIVE probe after 48 hours, and observed using a confocal microscope with fluorescence capabilities. The LIVE probe distinguishes live cells by measuring intracellular esterase activity. The non- fluorescent calcein, present in the LIVE probe, enters the cell membrane and is converted by enzymes in living cells to fluorescent calcein. The pictures shown in Figure 7 show that the hydrogel samples fabricated in SL have an even distribution of live cells. The pictures were taken at low (2.5X) magnification, and show approximately half of the cell-containing scaffold.Figure 7.

The pictures were taken at low (2.5X) magnification, and show approximately half of the cell-containing scaffold.Figure 7. Live stain of the cell-containing scaffold after 48 hr: visible light picture (left), fluorescent light picture (middle), and visible and fluorescent light pictures overlapped (right). These pictures were taken at low (2.5X) magnification.Figure 8 shows 10X magnification pictures of scaffolds fabricated at the three different conditions: single layer scaffold (1-mm thick) in the top row, single layer scaffold with laser pattern drawn twice (1-mm thick) in the middle row, and double layer scaffold (2-mm thick) in the bottom row. Qualitatively, increased numbers of live cells are observed in the single and double layer scaffolds (top and bottom rows) compared with the single layer scaffold fabricated with twice the UV laser energy.

The decreased number of cells in the successive exposure gel could be due to the increased UV exposure (at the laser wavelength of 325 nm). Work is ongoing to test the effect of successive UV exposure as well as other SL fabrication parameters on cell viability.442Figure 8. Live stain (after 48 hours) of the cell-containing scaffolds fabricated in SL. Top row: single layer scaffold (1-mm thick); middle row: single layer scaffold with laser pattern drawn twice; bottom row: double layer scaffold (2-mm thick). Left column: visible light pictures, middle column: fluorescent light pictures, and right column: visible and fluorescent light pictures overlapped.4. Conclusions and Future WorkThe use of SL for fabricating complex 3D tissue engineered PEG scaffolds was presented.The photopolymerization reaction of the materials used in SL should be sufficiently fast to produce thin polymer networks with adequate mechanical strength to be handled.

A practical method was described which determined the cure depth, or gel thickness, of a photoreactive hydrogel solution. The gel thickness curves obtained with this approach were utilized to successfully fabricate PEG-based scaffolds using SL. Gel thickness can be controlled by varying443the PI concentration or the energy dosage (using SL scan parameters and laser energy) for a given PI concentration. Curves of hydrogel thickness were utilized to successfully plan, perform, and demonstrate layered manufacturing of highly complex hydrogel scaffold structures, including structures with internal channels of various orientations. Successful fabrication of 3D, multi-layered bioactive PEG scaffolds containing cells was accomplished using a slightly modified commercial SL system and procedure.

Successful fabrication of 3D, multi-layered bioactive PEG scaffolds containing cells was accomplished using a slightly modified commercial SL system and procedure. Human dermal fibroblast (HDF) cells were encapsulated in bioactive PEG hydrogels fabricated using SL and successfully demonstrated to survive the fabrication process.The combined use of SL and photocrosslinkable biomaterials such as PEG makes it possibleto fabricate complex 3D scaffolds that provide site-specific and tailored mechanical properties (using multiple polymer materials) with a polymer matrix that allows transport of nutrients and waste at the macroscale and facilitates cellular processes at the microscale through precisely placed bioactive agents. The preliminary results presented here demonstrate that SL is a capable hydrogel layered manufacturing technology with potential for fabricating TE scaffolds incorporating precisely placed bioactive agents and live cells.

Additional work is necessary to quantify cell viability in the fabricated scaffolds, and to determine what effects other SL factors (such as laser wavelength, successive energy exposure, radical formation during photocrosslinking, and others) have on cell viability and proliferation.AcknowledgmentsThis work was funded, in part, through Grant Number 11804 from the W.M. KeckFoundation and the Texas Advanced Research (Advanced Technology/Technology Development and Transfer) Program under Grant Number 003661-0020-2003. KA was supported, in part, through Grant Number 0245071 from the National Science Foundation. Research support was also provided through the endowed Mr. and Mrs. MacIntosh Murchison Chair I in Engineering. The technical assistance provided by Ms. Gladys Almodovar, Ms. Lorena Le Maitre, Ms. Marian Viveros, Dr. Kristine Garza, and Dr. Joanne Ellzey from the UTEP Department of Biological Sciences was greatly appreciated.

Gladys Almodovar, Ms. Lorena Le Maitre, Ms. Marian Viveros, Dr. Kristine Garza, and Dr. Joanne Ellzey from the UTEP Department of Biological Sciences was greatly appreciated. Equipment and facilities in the Biological Sciences Department used here are maintained through NIH Grant Number G12 RR08124 from Research Centers at Minority Institutions (RCMI), NIH National Center for Research Resources. Special thanks to Dr. Kristine Garza for her advice and providing the LIVE/DEAD® assay presented in this work.ReferencesAng, T.H., Sultana, F.S.A., Hutmacher, D.W., Wong, Y.S., Fuh, J.Y.H., Mo, X.M., Loh, H.T., Burdet, E., and Teoh, S.H., “Fabrication of 3D Chitosan-Hydroxyapatite Scaffolds Using a Robotic Dispensing System,” Materials Science and Engineering C, 20:35-42, 2002. Arcaute, K., Ochoa, L., Medina, F., Elkins, C., Mann, B.

Bredesen Center, University of Tennessee-Knoxville, Knoxville, TN 37996 4. Materials Science and Technology Division, ORNL, Oak Ridge, TN 37831 5. Energy and Transportation Science Division, ORNL, Oak Ridge, TN 37831Corresponding Author: Zachary Cordero; zc29@rice.edu; 713-348-2495Abstract: We have examined the hypervelocity impact response of targets made from monolithic A356 and 316L stainless steel, as well as an additively manufactured A356/316L interpenetrating phase composite. 1.9 mm diameter spherical projectiles made from 2017 aluminum were fired at velocities of 5.9-6.1 km/s, allowing for the observation of multiple types of macro- and microstructural damage within each target. The macroscopic cratering damage to the A356/316L composite resembles that of the A356, but observations of both the cross section and the microstructural damage suggest that the A356/316L composite may be more resistant to spalling than A356 shielding with the same areal density.1.

Introduction In a recent study, we described a two-step approach for synthesizing net-shaped metallic composites [1]. In the first step of this approach, selective laser melting is used to fabricate a lattice preform that serves as the reinforcement. In the second step, this lattice preform is liquid metal infiltrated with a second material that has a melting point lower than that of the lattice. The resulting part is an interpenetrating phase composite (IPC) in which both constituents form continuous networks. This print-and-infiltrate approach offers two key advantages over other techniques for synthesizing metallic composites. First, because this approach excludes liquid- phase mixing of the constituents, it largely avoids the problems with intermetallic formation that limit other fusion-based techniques for printing metallic composites [2].

Second, it can pattern the constituents with a resolution on the order of 100 µm, or at least an order of magnitude finer than that of fusion-based techniques that require changing the feedstock composition during printing [3]. In our preliminary studies, we have used this two-step processing route to fabricate proof-of- concept composites of 316L stainless steel and a heat-treatable aluminum casting alloy, A356. In this case the 316L is the reinforcement and the A356 is the infiltrant. These first-generation A356/316L composites can have low densities (<4.10 g/cm3), respectable yield strengths (140-250 MPa), higher thermal conductivities compared to 316L (44 W/m K), and exceptional damage tolerance under tensile loading.

For example, tension tests on composites containing 38 vol% 316L1 2341showed a strain-to-failure of 32%, representing an order of magnitude improvement over the strain-to-failure of monolithic A356 [1].This preliminary work demonstrated that these additively manufactured composites exhibit a suite of properties that makes them especially attractive candidates for lightweight armor used in multiple-hit aerospace applications. Motivated by these findings, the objective of this study was to evaluate the hypervelocity impact response of these additively manufactured A356/316L IPCs and to compare their performance with those of the constituent materials. Hypervelocity impacts are of special importance in aerospace shielding applications, where shielding materials are routinely impacted by dust and debris orbiting the Earth at speeds of over 5 km/s [4].2. Experimental Methods 2.1.

Experimental Methods 2.1. Materials Processing We evaluated the hypervelocity impact response of an additively manufactured 38 vol% 316L composite prepared using the method described in [1]. First, we printed the 316L lattice preforms using a Renishaw AM250 selective laser melting system, using a feedstock of gas-atomized 316L powder (Fe-18Cr-12-Ni-2Mo, wt%). The lattice had body centered cubic symmetry, the unit cell geometry shown with schematically in Figure 1. The unit cell edge-length was 2.5 mm, and the 316L strut diameter was 0.8 mm. Next, we used centrifugal casting to infiltrate the lattice preforms with A356. Prior to impact testing, the strike face and rear surface of the composite targets were face-milled and polished to mirror finish. For comparison purposes, we also performed impact tests on cold-rolled 316L and as-cast A356. The final dimensions of the 316L, A356, and composite targets were 41x40x12.5 mm3, 40x39x12 mm3, and 38x36x10.9 mm3, respectively.

The final dimensions of the 316L, A356, and composite targets were 41x40x12.5 mm3, 40x39x12 mm3, and 38x36x10.9 mm3, respectively. Table 1 summarizes the densities of the cold-rolled 316L, as-cast A356, and A356/316L IPC. Table 1 also includes the Vickers microhardness values for the cold-rolled 316L, as-cast A356, and the constituents of the IPC. The microhardness tests were performed under a load of 300 gf. Areal Density (g/cm2) 9.96 – 3.24 – 5.142 23422.2. Hypervelocity Impact Setup We used hypervelocity impact tests to compare the impact performance of the A356/316L composite to the performance of its constituents, A356 and 316L. We conducted the hypervelocity impact tests using an 0.17 caliber two-stage light-gas gun, which has the capability to propel small projectiles at speeds of up to 7 km/s. The two-stage light-gas gun achieves these velocities by first detonating a charge of smokeless gunpowder inside a pump tube, expanding against a piston.

The two-stage light-gas gun achieves these velocities by first detonating a charge of smokeless gunpowder inside a pump tube, expanding against a piston. This then forces the piston into a volume of hydrogen gas, bringing that gas under immense pressures, on the order of 10-1 GPa [5]. The compressed hydrogen accelerates the projectile and its nylon sabot down the length of the barrel. The sabot discards from the projectile as it leaves the barrel, and the projectile interrupts multiple laser/photodiode intervalometers, allowing the velocity of the projectile to be recorded.Each test used a 1.9 mm diameter sphere of 2017 aluminum as a projectile, which will produce impacts representative of those caused by micrometeriods and orbital debris [4]. The measured velocities and calculated projectile kinetic energies are summarized in Table 2. The projectiles struck the targets at normal incidence.Table 2: Velocity and kinetic energy of aluminum projectile.

The projectiles struck the targets at normal incidence.Table 2: Velocity and kinetic energy of aluminum projectile. Kinetic Energy (J) Velocity (km/s) 180 6.0 176 5.9 186 6.1316L A356 38 vol% 316L Composite2.3. Impact Characterization: After each impact test, we removed the gunpowder residue from the samples in an ultrasonic cleaning bath containing isopropanol. We performed optical profilometry on the craters using a Keyence VR-3000 Wide-Area 3D measurement system. Following the profilometry, we performed postmortem analysis on the impact targets using a FEI Quanta 400 SEM and a JEOL JSM-6560 SEM. Finally, we cross-sectioned, mounted, and polished the targets, and then examined the crater using optical microscopy. The 316L cross-sections were etched in 5% Nital to reveal grain structure.3. Results and Discussion After each material underwent impact testing, the three targets were characterized using standard metallographic procedures.

Results and Discussion After each material underwent impact testing, the three targets were characterized using standard metallographic procedures. Crater dimensions provide a simple metric for comparing the amount of damage each material sustains after it undergoes hypervelocity impact. Typical hypervelocity impacts display hemispherical, symmetric craters with a depth roughly half of their width [4]. The dimensions of the craters in the different target materials are listed in Table 3. The morphology of the craters and the microstructure in the impact zone are described in more detail below. Evidence of localized melting is present in every crater, as the contact temperatures during impact can easily exceed 1000 K [6].3 2343Table 3: Macroscopic target damage data316L A356 38 vol% 316L CompositeCrater Volume (mm3) 25 117 96Crater Depth (mm) 2.0 4.5 4.4Crater Width (mm) 5.1 7.2 6.83.1.

Damage Observed in 316L Stainless Steel Figure 2 is a low magnification SEM image of the 316L target’s crater. The hemispherical, symmetric crater shows a petalled lip typical of hypervelocity impact into ductile materials [7]. There was no fragmented material surrounding the crater, and no cracks were present outside of the crater. The crater depth was 1.05x larger than the projectile diameter, and the crater width was 2.7x the projectile diameter. Overall, there was only localized macroscopic damage on the 316L impact surface. Profilometry showed a 16 μm tall bulge protruding from the rear surface.Higher magnification SEM images revealed some of the damage mechanisms in the 316L. Slip traces with spacings on the order of 0.1- 5 μm were visible in a zone 0.5 mm around the petalled lip of the crater. Inside the crater, a uniform layer of aluminum with a 1 µm grain size was deposited as a result of projectile vaporization.

Inside the crater, a uniform layer of aluminum with a 1 µm grain size was deposited as a result of projectile vaporization. The the crater displayed many small steps, similar to those identified by Shockey et al., who suggested that these steps result from shear bands that terminate on the inner surface of the crater [8].inside ofFigure 2: Low magnification SEM image of damage to the front surface of 316L targetA cross-section micrograph of the polished and etched 316L is shown in Figure 3. There was no evidence of internal cracking. A small zone of plastic deformation was revealed, extending roughly 0.5 mm beyond the crater’s face, identical to the width of the zone displaying slip traces outside the crater. No spalling or cracking was observed on the back face of this target.3.2. Damage Observed in As-cast A356 Figure 4 is a low magnification SEM image of the crater in the A356 target. Inspection of the A356 shows that it sustained significantly more damage than the 316L.

Damage Observed in As-cast A356 Figure 4 is a low magnification SEM image of the crater in the A356 target. Inspection of the A356 shows that it sustained significantly more damage than the 316L. The hemispherical crater is 3.8x wider and 2.4x deeper than the projectile. Additionally, the crater volume of the A356 is 4.7x that of the 316L target’s crater. There are cracks visible on the surface, and the petalled lip that surrounded the crater in the 316L is absent in the A356. This lip is not present in the A356 target because the material ejected from the crater fractured as a result of the extensive plastic strains generated on impact.4 2344Visual inspection of the rear surface of the A356 target also showed extensive damage, due to incipient spall, creating a cracked dimple that extends a distance of 0.7 mm from the rear surface. A compressive shockwave is created as the projectile impinges upon the target, which travels towards the rear face of the target.

A compressive shockwave is created as the projectile impinges upon the target, which travels towards the rear face of the target. When this compressive wave reaches the rear surface, it reflects as a tensile wave that travels back towards the strike face. Spall occurs when the magnitude of the tensile stress in this reflected shock wave exceeds the ultimate dynamic tensile strength of the material [9]. For impacts between spherical projectiles and flat targets, this spalling criterion is met along a spherical section projected outwards from the crater, with a distance from the rear surface of 10-50% of the thickness of the plate [4]. Spall is capable of causing extensive damage, even if the target is not perforated by the projectile, and is a major concern in both spaceflight and military armor applications. In this test, material did not detach from the back face, but the large, bulged region, approximately 9.4 mm in diameter, is filled with extensive cracks.

In this test, material did not detach from the back face, but the large, bulged region, approximately 9.4 mm in diameter, is filled with extensive cracks. This suggests that the spall carried almost enough momentum to cause complete detachment from the rear surface of the target.Figure 3: Optical image of the cross-sectioned, polished and etched 316L target.Postmortem analysis conducted on the A356 target showed damage mechanisms similar to those found in quasi-static testing of A356, with cracks propagating along the brittle silicon eutectic particles in the interdendritic regions [10]. In the annular region surrounding the crater, the material showed cleavage facets, as cracks propagated between primary aluminum grains, then through interdendritic regions rich in brittle silicon particles. Small cracks were clearly visible on the strike face of the target, and extended at least 2 mm beyond the edge of the fractured individual lip.

Small cracks were clearly visible on the strike face of the target, and extended at least 2 mm beyond the edge of the fractured individual lip. Within aluminum grains in this region, slip traces with ~2 μm spacing were commonly observed. The interior of the A356 crater was slightly less uniform than the interior of the 316L crater, with deposits of projectile material forming small masses at the bottom of the crater. Towards the top of the crater there wastheFigure 4: Low magnification SEM image of damage to the front surface observed on A356 target.5 2345evidence of brittle fracture identical to that of the fractured region at the top of the crater. On the back face of the specimen, the cracks propagated through the interdendritic regions. The damage modes seen in our specimens are broadly consistent with those reported by Brewer et al. [11] and Wang et al.

The damage modes seen in our specimens are broadly consistent with those reported by Brewer et al. [11] and Wang et al. [12] in their respective studies on impact into aluminum and aluminum alloys.Analysis of the A356 target’s cross section, in Figure 5, revealed extensive shown microcracking the distributed thickness of the target. The incipient spall plane was 1.5 mm from the back face and was approximately 1 cm in diameter. Many cracks were present, the largest of which were ~2 mm long. The largest cracks were oriented parallel to the strike face but there were also many small cracks that were perpendicular to the strike face and that traveled through interdendritic regions between the primary aluminum grains. A region of intense plastic deformation, where the dendrites were extremely compressed, extended 1 mm away from the crater wall.

A region of intense plastic deformation, where the dendrites were extremely compressed, extended 1 mm away from the crater wall. These observations show that the A356 was severely damaged by the impact, and that it would perform poorly in multiple-hit applications.throughFigure 5: Optical image of the cross-sectioned and polished A356 target.3.3. Damage Observed in 38 vol% 316L Composite The surface of the A356/316L composite showed a variety of damage modes, reflecting the disparate impact behaviors of the two component materials. The SEM image of the crater in Figure 6 shows that the crater was square in shape rather than circular, reflecting the symmetry of the lattice. There was neither a large raised lip, as in the 316L target, nor a significant fractured region, as in the A356 target. The square crater was 3.6x wider and 2.3x deeper than the projectile, or approximately 3 unit cells wide and 1 unit cell deep.

The square crater was 3.6x wider and 2.3x deeper than the projectile, or approximately 3 unit cells wide and 1 unit cell deep. The A356 at the edge of the crater eroded 1 mm further than the steel, as the steel shielded the remainder of the A356 from damage. Profilometry on the rear surface of the target revealed that the aluminum was slightly extruded through the steel, and that the bulge on the composite was only slightly higher than the bulge on the rear surface of the 316L target, which was also 1.6 mm thicker than the composite.Figure 6: Low magnification SEM image of damage to front surface of the A356/316L target.6 2346Close inspection of the composite revealed that it shared some traits in common with the 316L and A356 targets, but that the crater in the composite was not nearly as uniform as the craters in the two monolithic targets.

The A356 in the composite appeared to have fractured at the edge of the crater, displaying crevices approximately 0.5 mm deep between the 316L struts. This region showed the same fracture mechanisms as the A356: a combination of brittle cleavage and intergranular fracture between the primary aluminum grains. Likewise, the steel showed evidence of plastic deformation in the struts surrounding the crater, with similarly spaced slip traces present in both the printed and as-received 316L. No surface cracking was present in the aluminum outside the eroded and plastically deformed region, and the only cracks present in the steel were localized to small cracks between pores in the single unit cell adjacent to the crater edge.

The tips of the upturned struts in the composite specimen resemble the petalled lips of the monolithic 316L crater, with fine ductile dimples found on the surface where failure occurred.The cross-section micrograph of the A356/316L composite shown in Figure 7 reveals some of the unique damage modes in these IPC shielding elements. One interesting feature is the morphology of the 316L struts. The struts deformed upon impact, spreading apart from the center of the unit cell, and forming a protective coating over the A356. This deformation and protection occurred as a direct result of the crater’s rapid expansion. As material was driven out of the crater, the aluminum ruptured and was ejected, leaving behind the angled steel struts, which were oriented so that when they deformed with the expanding crater, they bent outwards from the center, increasing the local areal density and resistance to further deformation, halting the expansion of the crater.

Figure 7 shows that plastic deformation was limited to the unit cell adjacent to the impact, and that there was no cracking present in either the 316L or the A356. In addition, there was no evidence of spall in the composite material under this set of test parameters. We hypothesize that this may be because the compressive shockwave generated on impact is attenuated as it travels through the composite, as it closes pores and other residual defects. Because there is no damage evident in the region outside of the immediate vicinity of the crater, these results demonstrate the potential for this material to withstand multiple hits.Figure 7: Optical image of the cross-sectioned and polished A356/316L composite target.4. Summary We have developed an additively manufactured A356/316L IPC that may offer improved shielding performance for air- and spacecraft when compared to either A356 or 316L monolithic armor, by remaining lightweight while localizing damage to a small area.

This composite has demonstrated increased spall-resistance when compared to monolithic A356, and microstructural damage does not appear to occur outside of the immediate vicinity of the crater. In many different metrics of 7 2347shielding performance, this material behaves well, due to its unique traits resulting in the composite efficiently dissipating the kinetic energy of the projectile.AcknowledgementsThe authors wish to thank R. J. McCandless of Jacobs Technology for valuable assistance in operating the light-gas gun, and T. Seabury of ExxonMobil for helping with the metallographic investigations. This project was supported by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Vehicle Technologies Office, as part of the Propulsion Materials Program.This manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05- 00OR22725 with the U.S. Department of Energy.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceILLUMINATION AND IMAGE PROCESSING FOR REAL-TIME CONTROL OF DIRECTED ENERGY DEPOSITION ADDITIVE MANUFACTURINGD. Seltzer*, J. L. Schiano*, A. R. Nassar†, and E. W. Reutzel†School of Electrical Engineering and Computer Science, The Pennsylvania State University, University Park, PA 16802 †Applied Research Laboratory at The Pennsylvania State University, University Park, PA 16802AbstractThis paper describes the optical setup and image processing required to estimate melt-pool width and build height for real-time control of melt-pool geometry in directed energy deposition additive manufacturing. To overcome optical interference from plasma emissions and laser interactions, the melt-zone is imaged using laser illumination.

To overcome optical interference from plasma emissions and laser interactions, the melt-zone is imaged using laser illumination. A single camera, fixed to the processing laser, views the laser interaction zone and provides images for estimating melt-zone width and build height. Using a bandpass filter and a single aspheric lens, the camera system provides sufficient magnification and depth of field to achieve a 1-mil (25.4 µm) resolution. Maintaining melt-zone geometry within desired tolerances requires an image acquisition and processing rate on the order of 100 frames per second. This bandwidth is achieved by a Camera Link camera and field- programmable gate array that implements algorithms for estimating melt-pool width and build height. The design and experimental verification of the camera, illumination, and processing systems are discussed.1.

The design and experimental verification of the camera, illumination, and processing systems are discussed.1. IntroductionIn order to provide vision-based feedback control, which is essential for maintaining part quality in metal-based additive manufacturing (AM) processes, images of melt-pool geometry must be both acquired and processed in real-time. Based on previous studies [1], in order to obtain images without interference from black body and plasma-plume emissions caused by laser interaction it is necessary to illuminate and image the melt-pool. For this work we used light at a 405 nm wavelength in order to take advantage of the lower plasma-plume emission intensity at the given wavelength.

For this work we used light at a 405 nm wavelength in order to take advantage of the lower plasma-plume emission intensity at the given wavelength. Melt-pool geometry measurements obtained in real-time from this vision system are then verified using post-deposition measurements with an optical profilometer.The aim of this work is to design an imaging system that can determine build height and melt-pool width for later use in a vision-based feedback control system. The intended final goal being a real- time feedback control system that regulates build height and melt-pool width by controlling scanning speed and laser power inputs to the system.Two results can be concluded from this work. First, 405 nm is a viable wavelength for illuminating and imaging the melt-pool during the AM process. Second, it is necessary that the imaging system has both sufficient illumination and speed for obtaining real-time measurements.14792.

Second, it is necessary that the imaging system has both sufficient illumination and speed for obtaining real-time measurements.14792. Experimental SetupThe AM processing experiments used an Optomec LENS MR-7 directed-energy-deposition system. The LENS system utilizes a 500 W, Ytterbium-doped fiber-laser (IPG YLR-500-SM). The working distance is 0.365 in, measured from the substrate to four radially-symmetric power delivery nozzles. Centered within the nozzles is a center-purge nozzle through which argon flows coaxially onto the substrate.A custom mounting fixture, shown in Figure 1, extends from the laser processing head. Mounted on this fixture is a monochrome camera and a 405 nm illumination source, both aimed at the front of the melt-pool such that the melt-pool appears to be moving towards the camera. The optical axis of the camera is approximately 15° with respect to the substrate. An additional custom mounting fixture surrounds the laser processing head.

The optical axis of the camera is approximately 15° with respect to the substrate. An additional custom mounting fixture surrounds the laser processing head. Mounted on this fixture are four radially- symmetric 405nm illumination sources.Figure 1. Camera and initial illumination system mounting fixture.The camera used is an ACER model acA2040-180km Camera Link camera. Using an 8-tap Camera Link configuration allows for transmission of a 2048 x 2048, 8-bit monochromatic image in under 1ms. This high-speed transfer allows for a maximum frame rate of up to 180 frames per second (fps). Using a single aspheric lens with a 405 nm bandpass filter, configured to provide approximately 2.5X magnification, yields a resolution of 0.08 mils per pixel.The illumination source attached to the camera mount is a 170 mW, 405 nm diode-laser. A laser speckle reducer, model LSR-3005-6D-VIS, with a focusing lens were used with this laser.

A laser speckle reducer, model LSR-3005-6D-VIS, with a focusing lens were used with this laser. This1480configuration was only used during initial experiments. The radial illumination sources consisted of four 900 mW, 405 nm diode-lasers. With respect to the melt-pool and the camera at 0°, these lasers were positioned at ± 45° and ± 135° around the melt-pool. No laser speckle reducer or focusing lenses were used for these lasers.A dedicated PXIe chassis, with a field-programmable gate array (FPGA) and Camera Link adapter module, acquired camera images and system parameters such as laser power, scan speed and position.The data used in this work was obtained in two sets of experiments. In both sets, scanning speed and laser power undergo step changes from their nominal values, 25 ipm and 350 W respectively, as shown in Figure 2. The first set of experiments utilizes only the 750 mW with laser speckle reduction illumination configuration.

The first set of experiments utilizes only the 750 mW with laser speckle reduction illumination configuration. While the second set of experiments use only the four radial without laser speckle reduction configuration.Figure 2. Laser Power and Scan Speed input commands for experiments.For each experiment data is acquired every 50 ms. Because the camera is mounted to the processing head, the melt-pool remains fixed with respect to the horizontal image plane. This fact allows the image processing algorithm to search along a given vertical vector in order to determine the center of the melt-pool. Once the center of the melt-pool is located, build height and melt-pool width are estimated using threshold detection to determine the boundary between the melt-pool and solidified metal.The algorithm for determining melt-pool width and build height is tested by manual selection of melt-pool boundaries for each image obtained.

Optical profilometry is then used following the experiments to establish a ground truth for melt-pool width and build height. The build height and melt-pool width are determined approximately every 0.0002 in along the deposit, and because scan speed and position are recorded throughout each experiment the optical profilometer data can be1481plotted as a function of time or space. This allows for comparison of temporal camera data with spatial profilometer data.3. Results and Data AnalysisMelt-pool width and build height derived from the camera images follow the trends observed in the optical profilometer data. The camera images show a boundary between the melt-pool and solidified metal. However, insufficient illumination results in significant error in the algorithm’s estimation of melt-pool geometry.

The camera images show a boundary between the melt-pool and solidified metal. However, insufficient illumination results in significant error in the algorithm’s estimation of melt-pool geometry. In the first set of experiments the algorithm was unable to find any of the melt-pool boundaries, and manual determination while possible was often extremely difficult as demonstrated by figure 3. Additionally, not all plasma-plume emissions were able to be removed at this level of illumination.Figure 3. Camera image at time 2.5 seconds during first set of experiments.In the second set of experiments, the additional illumination provided by the array of radial laser sources allowed the algorithm to determine melt-pool boundaries as demonstrated by figure 4 and figure 5. In addition, unlike the first set of experiments, this level of illumination is sufficient to completely remove interference caused by plasma-plume emissions.1482Figure 4.

In addition, unlike the first set of experiments, this level of illumination is sufficient to completely remove interference caused by plasma-plume emissions.1482Figure 4. Camera image at time 2.5 seconds during second set of experiments.Figure 5. Camera image at time 2.5 seconds during second set of experiments with build height and melt-pool width parameter box drawn as estimated by melt-pool geometry algorithm.1483A comparison between the algorithm’s boundaries and manually selected boundaries reveals that the illumination was insufficient to provide an acceptable signal-to-noise ratio, resulting in significant error of melt-pool geometry estimates. Figures 6 and 7 show a closer look at relative pixel intensities along the vertical and horizontal search vectors used for estimating build height and melt-pool width respectively. The vertical search vector is a constant vector along the processing laser’s position selected through calibration images prior to processing.

The vertical search vector is a constant vector along the processing laser’s position selected through calibration images prior to processing. The horizontal search vector is selected by the algorithm during processing based on the estimated build height and boundary positions on the camera image. Poorly defined boundary points results in estimation error by the boundary algorithm.Figure 6. Pixel intensities along build height search vector over time (left). Examples of a well- defined set of height boundaries (top-right) compared to a set of poorly-defined boundaries (bottom-right) for the same scan speed and laser power settings.1484Figure 7. Pixel intensities along width search vector over time (left).

Pixel intensities along width search vector over time (left). Examples of a well- defined set of width boundaries (top-right) compared to a set of poorly-defined boundaries (bottom-right) for the same scan speed and laser power settings.A comparison of both manual and algorithmic camera estimates to profilometer data shows in figures 8 and 9 that image measurements follow profilometer trends. However, this comparison also reveals that a sample rate of 50 ms is an insufficient sampling rate. The lack of sharp transitions for camera data following a step-change in a system input reveals that these expected output changes are occurring unobserved in between image samples. Because the end goal is to design a control system, it is desirable to have a sample rate such that there are at least 5 samples during a step-transition [2].Figure 8.

Because the end goal is to design a control system, it is desirable to have a sample rate such that there are at least 5 samples during a step-transition [2].Figure 8. Comparison graph of build height obtained from measurements of manual boundaries, algorithmic boundaries, and optical profilometer data.1485Figure 9. Comparison graph of melt-pool width obtained from measurements of manual boundaries, algorithmic boundaries, and optical profilometer data.4. ConclusionThe results of this study show that 405 nm is a viable wavelength to observe the AM process without interference from either black body or plasma-plume emissions. Our current configuration will require additional illumination for future experiments. By providing additional illumination, we can ensure that the melt-pool geometry algorithm can accurately and reliably estimate melt- pool boundaries.

By providing additional illumination, we can ensure that the melt-pool geometry algorithm can accurately and reliably estimate melt- pool boundaries. Additionally, providing more illumination will allow for a reduction in the camera’s exposure rate which was the primary limiting factor in determining the sample rate used during these experiments.Once the illumination system has been improved to provide the necessary amount of illumination for real-time observation of melt-pool geometry, system identification and process models can be created based on real-time images. This will allow for the design of a multi-input multi-output system that can regulate build height and melt-pool with using laser power and scan speed.AcknowledgementsThe authors would like to thank Ms. Dalonia O. Lenoir and Dr. Timothy B. Tighe for their assistance with the Optical Profilometry measurements. This work was supported by the Office of Naval Research, under Contract No. N00014-11-1-0668.

Lenoir and Dr. Timothy B. Tighe for their assistance with the Optical Profilometry measurements. This work was supported by the Office of Naval Research, under Contract No. N00014-11-1-0668. Any opinions, findings, and conclusions or recommendations expressed in this publication are those of the authors and do not reflect the views of the Office of Naval Research.References[1] D. Seltzer, X. Wang, A. Nassar, J. Schaino, and E. Reutzel, “System identification and feedback control for directed-energy metal-based additive manufacturing,” in Solid Freeform Fabrication Proceedings, Austin, TX. 2015.[2] Charles L. Phillips and H. Troy Nagel. Digital Control Systems Analysis and Design, thirdedition, Prentice Hall, 1995.1486.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperImage Registration and Matching Error in 2D and 3D for Laser Powder Bed FusionAndrew Lang1,3, Cesar Ortiz Rios2, Joseph Newkirk2, Robert G. Landers3, James Castle1, and Douglas A. Bristow31The Boeing Company 2Materials Science and Engineering, Missouri University of Science and Technology 3Mechanical and Aerospace Engineering, Missouri University of Science and TechnologyAbstractThis work outlines a method to register 2D and 3D images taken post-process and in situ from 301L stainless steel parts printed by Laser Powder Bed Fusion. The process uses DREAM.3D, an open-source software that provides for data transport in a non-proprietary format. The Robust Automatic Threshold selection technique is used to create a boundary point cloud of the part from each image.

The Robust Automatic Threshold selection technique is used to create a boundary point cloud of the part from each image. The Iterative Closest Point technique is applied to the point clouds for both 2D images and 3D image stacks to create an affine transformation matrix for registration. Multiple 2D SEM images of the same sampled layer are taken under different settings and imaging conditions and registered to a common target. Images from post-process X-ray Computed Tomography and an in situ short-wave infrared camera create 3D image stacks, which are directly registered in 3D space. Registration accuracy is validated by creating a correspondence list of the closest point in the registered point clouds and the matching error is calculated using mean average error. Mean average error is computed using point-to-point and point-to-plane methods; the point- to-plane method is shown to be more reliable.

Mean average error is computed using point-to-point and point-to-plane methods; the point- to-plane method is shown to be more reliable. Finally, the registered 3D images are down sampled to the lower resolution image dimensions creating arrays of in situ and post-process data at the same resolution.1.0 IntroductionImage registration is frequently used in Laser Powder Bed Fusion (LPBF) research to align multiple modalities of data, analyze variation in images, and fuse training datasets for classification. In registration, a reference dataset is transformed to align with a fixed target dataset [1]. There are two broad strategies for registration: 1) manual alignment [2, 3] and 2) automated algorithms, the most mainstream being Iterative Closest Point (ICP) [4].

There are two broad strategies for registration: 1) manual alignment [2, 3] and 2) automated algorithms, the most mainstream being Iterative Closest Point (ICP) [4]. Extensive research has been conducted on other registration algorithms that optimize a cost function to minimize error between two datasets using methods like defining datums from part features [5], quasi-Newton optimization [6, 7], least squares optimization [8], or mean squared difference [1].

Registration accuracy can be subjectively estimated by manually inspecting alignment [9] or objectively determined by calculating an error metric between points [8], but it is often not discussed in LPBF research.Fusing in situ data, often multiple modalities, with ex situ X-ray Computed Tomography (XCT) and/or destructive testing (DT) sectioned images has become a common area of research in quality assurance for metal additive manufacturing, both in industry and academia [10, 11, 12].1431This data fusion can be used for statistical analysis or machine learning models that predict resulting material characterization based in in situ parameters. There are many existing commercial tools for in situ monitoring of LPBF processes that are integrated into commercial systems, can be purchased as machine add-ons, or are independent of a particular machine concept [13]. XCT has become standard in industrial verification of additively manufactured (AM) parts [12].

XCT has become standard in industrial verification of additively manufactured (AM) parts [12]. XCT is advantageous for quality control because it is non-destructive and, depending on the resolution required, can relatively quickly capture profiles of a whole part. Micrographs from Scanning Electron Microscopes (SEM) can be used for microscopic material characterization but the part must be destroyed to use SEM and only samples of the part can be taken [14, 15].Many toolboxes have been developed to perform image registration. Some examples are Slicer, ImageJ, Elastix, Amira, DREAM.3D, and VGSTUDIO MAX. The Insight Toolkit (ITK) is an open-source development framework with an extensive suite of software tools for image analysis, including registration [16, 17], and is integrated in several analysis tools like DREAM.3D or Elastix.

Many of publicly available toolboxes and registration techniques were initially developed for medical imaging while LBPF imaging research was still in its infancy [18, 19]. There are several image processing libraries, like OpenCV, Pillow, and scikit-image for Python, MATLAB, and other languages that can be utilized for 2D image registration. The main disadvantages with most available toolboxes are that their application to specific needs of LPBF research is limited, they may not work with 3D image sets or triangle geometry, and they are not tightly integrated with a standard visualization tool.This paper investigates image registration in both 2D and 3D utilizing DREAM.3D, an open-source tool kit that allows for construction of customized workflows to analyze data [20]. DREAM.3D was developed to process digital instances of microstructure, but many of the tools are useful for LPBF data processing and address the disadvantages found in most available toolboxes.

DREAM.3D was developed to process digital instances of microstructure, but many of the tools are useful for LPBF data processing and address the disadvantages found in most available toolboxes. ParaView, an open-source analysis and visualization platform, is used to render 2D and 3D visualizations (paraview.org). The main advantages of DREAM.3D are that the reconstruction, manipulation, and visualization of 3D data is simple and straightforward, the program is free and open-source, over 100 filters from the image processing library ITK are included, data is stored as standard HDF5 files by default, additional features are continuously added, and custom filters can be created.DREAM.3D processes two broad types of data: 1) image geometry and 2) triangle geometry [21]. Image geometry is a structured grid of constant resolution with data values assigned to each cell. The pixel makeup of a computer screen is an example of image geometry.

Image geometry is a structured grid of constant resolution with data values assigned to each cell. The pixel makeup of a computer screen is an example of image geometry. Triangle geometry consists of vertex points, edges forming lines between points, and faces forming a mesh between lines (faces could be triangles, quads, etc). A geodesic dome is a physical example of triangle geometry. Figure 1 shows a 2D example of a triangle geometry structure (a) with vertices, edges, and faces and an image geometry structure (b) with grayscale values in each cell of the grid.1432Figure 1: Triangle Geometry (a), Image Geometry (b)Datasets from many LPBF in situ and ex situ platforms are often structured voxel/pixel- based image geometry [13, 12]. The datasets used in this study are structured image datasets in both 2D and 3D.

The datasets used in this study are structured image datasets in both 2D and 3D. Currently, the DREAM.3D filter Apply Transformation to Geometry cannot apply a pre-computed transformation matrix to image-type geometries, so a workaround has been implemented utilizing Python for 2D images and a not-fully-automated process involving manual rotation for 3D images.2.0 Experimental SetupFor this study, an ASTM E8 tensile specimen [22] has been printed in 304L stainless steel using a Renishaw AM250. The test specimen is 50 mm tall and has a 4 mm diameter neck. The nominal laser power for this build was 200W and a section of lettering (“Missouri S&T”) was printed at 100W inside the neck of the part to induce defects. The point distance, dp, and the hatch spacing, dh, were held at constant 60 μm and 85 μm, respectively.

The point distance, dp, and the hatch spacing, dh, were held at constant 60 μm and 85 μm, respectively. In-situ data was captured using a Short Wave Infrared (SWIR) camera, XCT imaging was used to generate a 3D array of the part density, and 2D SEM images were taken of extracted sections of the test specimen.Figure 2: ASTM E8 tensile specimen (a), low power lettering printed in the neck to induce defects (b), the final tensile bar printed by LPBF (c), 3D reconstruction of in situ SWIR imaging (d), and 3D reconstruction of XCT imaging (e).14332.1 SWIR ImagingThe purpose of SWIR imaging for this study was to establish a 3D profile of an in situ dataset to register with an ex situ dataset. The SWIR camera set up and thermal feature extraction has been performed in other work [23]; the process is summarized in this section. An FLIR SC6201 SWIR camera measured visible light emitted from the melt pool.

An FLIR SC6201 SWIR camera measured visible light emitted from the melt pool. The camera was installed at an observation angle of 15⁰ above the build chamber to observe the build. The camera pixel array was 2500 Hz). The x- reduced to an 80 × 80 pixel window enabling high frame rate recording ( direction instantaneous field of view of the SWIR camera was 130 μm/pixel. The y-direction instantaneous field of view was 135 μm/pixel due to the observation angle (θ = 15°) of the SWIR ∼ camera shown in Figure 3. A non-uniformity correction (NUC) was performed to account for differences in the SWIR measurements across the imaging area due to the observation angle. The final corrected pixel dimensions were 125 μm x 125 μm for each layer.∼∼Figure 3: Schematic of SWIR camera observations [23]Thermal features from a multi-frame time-series recording are extracted to a single image representation of each layer.

Each 2D layer is concatenated to create a 3D voxel-based reconstruction of the tensile specimen in HDF5 format. The layer height of this build was 50μm. An image geometry container was created in DREAM.3D with the X, Y, and Z dimensions of the voxels (125 μm, 125 μm, 50 μm) and the HDF5 thermal feature file was imported using the Import HDF5 Dataset filter and can be visualized in ParaView.2.2 X-Ray Computed Tomography ImagingPost-process XCT inspection was performed by NSI using a custom radiograph tool. NSI scanned the test piece with a MeVX6™ High Energy system using a 6 MeV linear accelerator (LINAC) at 450 kV and performed 360⁰ step scan imaging with a focal spot size of 24.15 μm. The full geometry of the tensile piece was captured in 291 x 281 x 1706 voxels. Each voxel is a cube with dimensions 29.96 μm in X, Y, and Z.

The full geometry of the tensile piece was captured in 291 x 281 x 1706 voxels. Each voxel is a cube with dimensions 29.96 μm in X, Y, and Z. The XCT array was directly imported into DREAM.3D using the Import North Star Imaging CT (.sihdr/.nsidat) filter.14342.3 SEM ImagingThe tensile bar was cut perpendicular to the long axis via wire EDM as shown in the cut plan in Figure 4 (a). The “M” section in the text printed at lower power was extracted between 36.65 and 31.65 mm and mounted in conductive mounting resin then metallographically prepared to a final polish of 0.05 µm for SEM analysis. During the metallographic preparation, 331 μm of material was removed, resulting in a cross section containing the targeted “M” low power region. Imaging parameters were selected to generate good contrast between the polished surface metal and the defect interior to facilitate registration in DREAM.3D.

Imaging parameters were selected to generate good contrast between the polished surface metal and the defect interior to facilitate registration in DREAM.3D. Images were captured with the Through-Lens Detector (TLD) in Backscatter Electron (BSE) mode biased to -150V voltage to eliminate the associated secondary electron (SE) edge effect. Integration of eight frames with a dwell time of 1 µs were used. The orientation of the 0⁰ imaging position was chosen to align to the orientation of the CT data. Images were taken in a 5x5 grid pattern starting from row 3, column 1 in increments of 1000 µm in the X-direction and 800 µm in the Y-direction. Increments were chosen to ensure sufficient overlap for stitching the images together. Once the images were captured, the stage was rotated to establish a new imaging position. The imaging process was repeated until the eight imaging positions, shown in Figure 4 (b), were captured.

Once the images were captured, the stage was rotated to establish a new imaging position. The imaging process was repeated until the eight imaging positions, shown in Figure 4 (b), were captured. An image manipulation program was used to stitch the images together. The images were imported as layers and overlayed using the overlapping features. Layers were ordered such that lower rows overlapped higher rows and interior columns overlapped exterior columns. For example, row 1 is overlapped by row 2, and column 3 overlaps both columns 2 and 4, which overlap columns 1 and 5, respectively. Pixel sizes of these images are 1.25 μm x 1.25 μm.When the images were arranged, the image was flattened and exported as a TIF file. The images were integrated into DREAM.3D using the ITK:Image Reader filter.

The images were integrated into DREAM.3D using the ITK:Image Reader filter. The Set Origin & Spacing (Image) filter was used to input the spacing based the pixel dimensions.Figure 4 Extracted section for SEM imaging: Cut Plan (a), 8 SEM imaging positions (b)3.0 RegistrationThis section outlines the processes used to register multiple modes of data in a single affine space. Boundary point clouds from each image were generated by sampling points along the surface of the part created using the Robust Automatic Threshold Selection (RATS) algorithm. An affine transformation was found to directly register two sets of point clouds using the ICP1435algorithm. Finally, DREAM.3D, Python, and ImageJ was used to transform the 2D and 3D datasets in a not-fully-automated process.3.1 Boundary Point CloudDREAM.3D was used to create point clouds representing the boundary of each 2D and 3D image array.

The Robust Automatic Threshold Selection (RATS) algorithm, an automatic thresholding method based on gradients in the image, was used to define the boundary of the part and point clouds were created by sampling from the boundary. RATS is an automated edge detector algorithm that identifies a threshold for greyscale images based on the image’s gradients, often at the maximum gradient [24, 25]. The gradient array was calculated using the ITK::Gradient Magnitude Image Filter. The magnitude of the image gradient is:𝐺𝐺(𝑥𝑥, 𝑦𝑦) = ‖𝑓𝑓(𝑥𝑥, 𝑦𝑦)‖2 = �𝑓𝑓𝑥𝑥(𝑥𝑥, 𝑦𝑦)The direction of the image gradient is:2+ 𝑓𝑓𝑦𝑦(𝑥𝑥, 𝑦𝑦)2−1 𝑓𝑓𝑥𝑥(𝑥𝑥, 𝑦𝑦) 𝑓𝑓𝑦𝑦(𝑥𝑥, 𝑦𝑦)𝜃𝜃(𝑥𝑥, 𝑦𝑦) = tanEliminating the square root of the magnitude yields suitable results without the added cost of a final scan across the image to compute the root.

In the ITK plugin, the gradient is simply the sum of the squares of the partial derivative operations.222= 𝑓𝑓𝑥𝑥(𝑥𝑥, 𝑦𝑦)+ 𝑓𝑓𝑦𝑦(𝑥𝑥, 𝑦𝑦)𝑔𝑔(𝑥𝑥, 𝑦𝑦) = 𝐺𝐺(𝑥𝑥,𝑦𝑦)With the Robust Automatic Threshold filter, DREAM.3D creates a new array that is false where the input array is less than the threshold and true otherwise. A regional threshold is computed as the gradient weight sum of the input array, A.𝑛𝑛 𝑖𝑖=1∑ �𝑔𝑔(𝑥𝑥, 𝑦𝑦) ∙ 𝐴𝐴(𝑥𝑥, 𝑦𝑦)� ∑ 𝑔𝑔(𝑥𝑥, 𝑦𝑦)𝑛𝑛 𝑖𝑖=1𝑇𝑇 =This threshold is generally where the gradient magnitude is highest.The ITK::Grayscale Fillhole Image filter, with the FullyConnected parameter set to true, is used to isolate and fill the part mask. To create a point cloud in the 2D images, the Find Boundary Cells (Image) filter is applied to the part mask and vertex geometry of each boundary cell is extracted to a point cloud.

To create a point cloud in the 2D images, the Find Boundary Cells (Image) filter is applied to the part mask and vertex geometry of each boundary cell is extracted to a point cloud. A visualization of the process is shown in Figure 5 for a 2D array using the 0⁰ TLD BSE SEM image.1436Figure 5: TLD BSE 0⁰ SEM image: Raw image (a), gradient magnitude array (b), part mask after applying RATS and filling holes (c), and boundary point cloud of the image (d).For both SWIR and SCT 3D datasets, triangle surface arrays were created from the surface of the mask array using the Quick Surface Mesh filter. Approximately four and a half million triangles were created for each surface. 3D point clouds were generated by sampling points from the triangle surfaces using the Point Sample Triangle Geometry filter. For registration, one hundred thousand sample points were chosen for both SWIR and XCT datasets.

For registration, one hundred thousand sample points were chosen for both SWIR and XCT datasets. A visualization of the process in 3D is shown in Figure 6 using the XCT array.3.2 Iterative Closest PointThe ICP algorithm iteratively revises a transformation to minimize the sum of squared differences between the coordinates of matched pairs and register two-point clouds to the same affine space [26]. There is a wide body of research to improve the algorithm by minimizing different error metrics, performing point-to-plane registration, or creating various search algorithms to reduce time to identify matched pairs [4, 27, 28]. This study used the DREAM.3D implementation of ICP to find an optimal affine transformation. The Iterative Closest Point filter finds a rigid transformation that rotates and transforms the reference data set.

The Iterative Closest Point filter finds a rigid transformation that rotates and transforms the reference data set. The filter does not change the distance between any two points in the transformed point cloud so there is no scale or shear component in the transformation.1437An affine transformation is a map between two different affine spaces that preserves lines and parallelism [25].

This implementation of ICP may get stuck at a local minimum so an initial guess of alignment was be provided before implementation. The initial guess for rotation was done using the Rotate Sample Reference Frame filter and translation was done using the Set Origin & Spacing (Image) filter. After the initial guess was applied, a transformation matrix was found with ICP and the reference point clouds were transformed in DREAM.3D.

After the initial guess was applied, a transformation matrix was found with ICP and the reference point clouds were transformed in DREAM.3D. Figures 7 and 8 show the 2D and 3D point cloud alignment before and after transformation.Figure 7: 2D point cloud registration: SEM with TLD BSE 0⁰ before (blue), after (red), and the target point cloud (white)1438Figure 8: 3D point cloud registration before/after: SWIR (white) and XCT (red) before (a), SWIR and XCT after registration (b and c)3.3 Apply Transformation to GeometryThe DREAM.3D filter, Apply Transformation to Geometry, can apply any 4x4 transformation to triangle geometry but, in this study, all three datasets (SEM, SWIR, XCT) are structured image geometries so the transformations computed in section 3.2 cannot be directly applied in DREAM.3D.

Registration for the 2D dataset relied on ImageJ and Python as a patch and registration for the 3D dataset relied on manual rotation before ICP.Unlike a point cloud, image geometry must be interpolated in the original grid size after it is transformed. Interpolation finds a value between two points on a line so, for this case, the interpolation algorithm determines the value to assign each voxel between two known points. In multidimensional interpolation, an estimate of y(x1, x2, …, xn) is made from an n-dimensional grid of tabulated values y and n one-dimensional vectors giving the tabulated values of each of the independent variables x1, x2, …, xn [29]. Bi-linear interpolation was used for all interpolations in this study.3.3.1 Apply Transformation to 2D GeometryThe eight SEM imaging positions are transformed with a combination of ImageJ, Python, and DREAM.3D. Before the images were imported into DREAM.3D they were all rotated to roughly align with the 0⁰ image position.

Before the images were imported into DREAM.3D they were all rotated to roughly align with the 0⁰ image position. As an example, Figure 9 (a) and (b) show the initial 60⁰ imaging position before and after the rotation. Bilinear interpolation was used, and the result was saved as a .tiff file which was could be read by DREAM.3D.Figure 9: TLD BSE 60⁰ SEM image: Raw image (a), image rotated to 0⁰ with ImageJ (b)Next, the ICP operation in DREAM.3D to determine a transformation matrix, as outlined in section 3.2, was performed. Python was used to apply the transformation to each of the eight1439SEM imaging positions. The DREAM.3D file was read into Python using the h5py library (h5py.org) and the applicable transformation matrix was read as a Numpy array (numpy.org). The image was then transformed with skimage.transform.warp using bilinear interpolation (scikit- image.org) and the final image was output as an .h5 file.

The image was then transformed with skimage.transform.warp using bilinear interpolation (scikit- image.org) and the final image was output as an .h5 file. The final image was read back into DREAM.3D and a surface point cloud was created in the final position.The rotation components of an affine transformation are:The translation components of an affine transformation are:Table 1 shows the transformation matrix, computed by ICP, for the 60⁰ imaging position. This transformation matrix indicates this is a 2D transformation as the third row and column are each [0, 0, 1, 0]. This transformation has a 2.16⁰ rotation about the Z-axis, as indicated by the four cells in the upper left quadrant of the matrix: cos-1(.999289) = sin-1(0.37712) = 2.16⁰ (Equation X). The translation components are tx = -0.067751 mm and ty = 0.08299 (Equation X). There are not scaling or shear components in this transformation matrix.

The translation components are tx = -0.067751 mm and ty = 0.08299 (Equation X). There are not scaling or shear components in this transformation matrix. All eight of the transformation matrices computed for the different SEM imaging positions follow the same format of 2D transformation, rotation about the Z-axis, and a small x and y translation.TLD BSE 60⁰ 0.999289 -0.0377521 0 00.0377521 0.999289 0 00 0 1 00.067751 0.08299 0 1Table 1: The transformation matrix computed by ICP for the TLD BSE 60⁰ SEM image. Dimensions are in mm.Figure 10 shows the initial location (a) and the transformed final location (b) of a section of the point clouds created from imaging positions. The improvement in registration is clear to see.

The improvement in registration is clear to see. A technique to measure the error in registration between the point clouds is outlined in Section 4.1440Figure 10: Initial location (a) and final location (b) of a section of the eight SEM imaging positions3.3.2 Apply Transformation to 3D GeometryAs discussed earlier, the Apply Transformation to Geometry filter in DREAM.3D does not work on image-type datasets. ImageJ and the Python libraries that were explored (scikit-image, Pillow, and OpenCV) only work on 2D images and cannot apply 4x4 transformations to 3D image- type datasets like what was done in Section 3.3.1 for the SEM imaging positions.3D images can, however, be manually translated and rotated in DREAM.3D using the Set Origin & Spacing (Image) and the Rotate Sample Reference Frame filter.

At iterative process where the SWIR array was manually rotated across all three axis, ICP was run to register the resulting point cloud to the XCT point cloud, and the computed transformation matrix compared to the translation-only matrix in equation X. This time-consuming process was stopped when a satisfactorily close solution to the translation-only matrix was achieved. Table 2 shows the resulting transformation matrix from this iterative manual-rotation/ICP process.SWIR 0.999995 -0.00523 0.000712 00.0052380 0.999988 -0.00063 00.00070846 0.000629 1.00001 00.0873794 -1.04344 0.472003 1Table 2: The transformation matrix computed by ICP for the SWIR image. Dimensions are in mm.The SWIR image array was translated to its final location with the Set Origin & Spacing (Image) filter using the translation components of the transformation, tx=0.08978 mm, ty=-1.04344 mm, and tz=0.472003 mm.

Figure 11 (a) shows the application of the translation components in DREAM.3D, (b) shows the SWIR array with the XCT array before the transformation, and (c) shows the SWIR array with the XCT array after the transformation.1441Figure 11: Translation components to register SWIR with SCT in DREAM.3D (a), SWIR array (red) before transformation to XCT (blue) (b), and SWIR array (red) after transformation to XCT (blue) (c).Applying only the translation components of the transformation in Table 2 will introduce some error in the final alignment because the rotation components are not accounted for. Some error, where the two edges are not integrated, is visible at the bottom (-Z) end of the neck of the tensile piece in Figure 11 (c).The registration accuracy can be interrogated by slicing the registered array in ParaView and examining internal or external reference features between the two arrays (Figure 12).

Internal features were imbedded into the tensile specimen by printing specific volume at lower laser power. Slices on the XY plane are shown in (a) at or three different reference features, slices on the YZ plane are shown in (b), and slices on the XZ plane are shown in (c). The internal features are visible in both the SWIR array and the XCT array. Both internal and external features were correlated across the SWIR and XCT arrays showing general alignment between the two datasets.Figure 12: Subjective verification of registration accuracy by comparing internal and external features in both arrays showing alignment along the Z axis (a), the X axis (b), and the Z axis (c). The slice number is shown with each image.14424.0 Registration AccuracyThis study found individual errors by calculating the Euclidian distance between matched points.

Where n is the number of points in the reference point cloud, M. The standard deviation, σ, for the errors was also calculated.Traditional ICP algorithms, including the one used in this study to find the transformation, attempt to minimize the Mean Square Error (MSE) of the distance between corresponding points in two point clouds [26, 30]. MSE, as a cost function in ICP algorithms, is advantageous over MAE because larger errors are penalized exponentially. MSE is in squared units which is a problem as an error metric because squared length units are not particularly useful in determining the distance error between point clouds. Root Mean Square Error (RMSE) is in the same unit length as the dataset so it reports a useable error metric.

Python was used to read point clouds from DREAM.3D and implement the error algorithm.A summary of the point-to-point error process:1. Create list of correspondence points, C.a. For all points in the reference point cloud, M, match closest point in targetpoint cloud, S.2. Calculate individual errors, ei.a. Find the Euclidian distance between matched pairs in M and C3. Calculate accuracy metricsa. Standard Deviation, MAE, MSE, RMSE144324.1.1 Point-to-Point Error with 2D ImagesPoint-to-point error was measured for each of the eight SEM imaging positions. Point clouds for the eight SEM imaging positions were the reference point clouds and the ninth human- labeled SEM image was the target point cloud. Table 3 below shows the accuracy metrics, standard deviation, MAE, MSE, RMSE, for the original and final point cloud positions. The table also shows the number of points, n, in each point cloud.

Table 3 below shows the accuracy metrics, standard deviation, MAE, MSE, RMSE, for the original and final point cloud positions. The table also shows the number of points, n, in each point cloud. MAE is improved in each point cloud using the registration technique from Section 3. The TLD 180⁰ imaging position has the closest final registration to the target, at MAE of 4.18 μm, and the 270⁰ imaging position has the worst final registration to the target, at MAE of 7.99 μm.

When the target point cloud or the target and the reference point cloud are randomly down sampled, the distance between matched points becomes larger. Error is unreliable if there are not a significant amount of points in the target point cloud. MAE diverges exponentially as the target and reference point cloud are down sampled as seen in Figure 13 (a) shown with a logarithmic scale. Figure 13 (b) shows the effect on MAE of down sampling only the reference point cloud. MAE is effectively unchanged as the reference point cloud is down sampled because the reference points are still able to find matched neighbors in the large target point cloud. As the number of points in the target are increased, the point cloud starts to act as a plane.

As the number of points in the target are increased, the point cloud starts to act as a plane. For a point-to-point error metric to be effective, it is vital to select a high number of points in the target point cloud.Figure 13: Change in MAE when the target point cloud is down sampled (a) and when the reference point cloud is down sampled (b)14444.1.2 Point-to-Point Error with 3D ImagesPoint-to-point MAE of the 3D data set, where the reference is the SWIR point cloud and the XCT is the target point cloud, was determined in the same way as the 2D SEM images in Section 4.1.1. Each point cloud had 100,000 points that were sampled off the surface as outlined in Section 3.1. Figure 14 shows a ParaView visualization of the two point clouds when the points are randomly down sampled to the same effect as the SEM images.Figure 14: The effect on MAE of down sampling both point clouds.

SWIR in red and XCT in white.When both point clouds are down sampled, MAE increases exponentially as seen in Figure 15 (a) shown with a logarithmic scale. Figure 15 (b) shows the effect of down sampling each point cloud separately. When the reference point cloud (SWIR) is down sampled the error effectively stays the same but when the target point cloud (XCT) is down sampled the error rises. The same effects were seen with the 2D SEM images. The point-to-point MAE for the 3D registration was 112 μm with 100,000 CT points and 86 μm with 1,000,000 CT points.1445Figure 15: Change in MAE when the target point cloud is down sampled (a) and when the reference point cloud is down sampled (b)4.2 Point-to-Plane ErrorICP registration of multiple images by point-to-plane has been developed to attempt to improve registration over traditional ICP [30].

The point-to-plane registration method is not used for this study, but the derivation of the error metric for point-to-plane registration has distinct advantages over the point-to-point error metric discussed in Section 4.1. The main advantage is the target is a triangle surface instead of a point cloud so point-to-plane does not depend on the number of points in the target. The second advantage is that DREAM.3D can calculate the individual point-to-plane errors using the Find Vertex to Triangle Distances which significantly simplifies the implementation of the algorithm and makes it more repeatable across platforms. Python is only used to calculate the array-level metrics (standard deviation, MAE, MSE, RMSE) which is a very straight forward calculation.Like point-to-point, the point-to-plane metric matches the closest triangles in the target surface for each point in the reference point cloud.

Individual errors are calculated using the distance between each source point and the tangent plane at its corresponding destination point [27, 30].Unit normals from the face of each triangle in the target surface are added to the triangle surface array in DREAM.3D with the Generate Triangle Normals filter. The error between the target surface and sampled points from the reference surface is shown in Figure 16 as ei.Figure 16: Point-to-plane error from a target surface and points on a source surface.1446A summary of the point-to-plane error process1. Read reference vertex points and the target triangle geometry surface in DREAM.3Da. Downsample the reference vertex points.2. Calculate individual errors, ei.a. For all points in the reference point cloud, M, calculate Euclidian distanceto the closest triangle in the target surface3. Calculate accuracy metricsa.

Calculate individual errors, ei.a. For all points in the reference point cloud, M, calculate Euclidian distanceto the closest triangle in the target surface3. Calculate accuracy metricsa. Standard Deviation, MAE, MSE, RMSE4.2.1 Point-to-Plane Error with 2D ImagesA triangle surface plane was created in DREAM.3D from the target SEM image. This was done by extrapolating the mask created by the RATS algorithm to several layers and creating a 3D array. The same process in Section 3.1 for 3D images was followed to create a triangle surface of the target image.

The same process in Section 3.1 for 3D images was followed to create a triangle surface of the target image. The mask is shown in Figure 17 (a) and the target surface with a point cloud from one of the SEM imaging positions (red) is shown in (b).Figure 17: An extruded target image mask (a) and a triangle boundary plane, based on the SEM target image, with one of the SEM imaging position point clouds (b).The triangle surface plane was the target and the boundary point clouds of each SEM imaging position were the reference point cloud. With this method, there was effectively no change in the MAE when the reference point clouds were down sampled. This is shown in Figure 18.

With this method, there was effectively no change in the MAE when the reference point clouds were down sampled. This is shown in Figure 18. Average MAE across all SEM imaging positions with no down sampling and with all down sampled results are both 2.4 μm.1447Figure 18: Point-to-plane MAE for SEM images with down sampling4.2.2 Point-to-Plane Error with 3D ImagesThe point-to-plane error for the 3D images used the triangle surface created for the XCT array in Section 3.1 to sample the boundary point cloud. The XCT triangle surface was the target and the SWIR point cloud was the reference. Like the 2D images, Figure 19 (a) shows the MAE stays effectively the same when the number of points in the SWIR point cloud were down sampled. The point-to-plane MAE for the 3D registration was, on average with five different levels of down sampling, 75 μm.Figure 19 (b) highlights the convergence of increasing the points in the target point cloud until it becomes a plane.

The scale is logarithmic, making the exponential trend toward the point- to-plane MAE results as the XCT point cloud is increased from 10 points to 1,000,000.Figure 19: Point-to-plane MAE for 3D images with down sampled SWIR point cloud (a) and comparing point-to-point and point-to-plane MAE5.0 Data FusionTo fuse each 3D dataset on a common grid, the DREAM.3D Fuse Regular Grids (Nearest Point) is used. The XCT array, along with the final position of the SWIR array, are fused together. The grid of cells in the reference data container (SWIR) is overlaid on the grid of cells in the sampling (XCT) data container. Each cell in the reference container is associated with the nearest point in the sampling container. The values of the nearest point in the sampling container are assigned to that cell in the reference container. This is a nearest neighbor operation as the values of the sampling voxels are not interpolated to the reference voxels.

This is a nearest neighbor operation as the values of the sampling voxels are not interpolated to the reference voxels. As a result, the sampling array,1448XCT in this case, is in the same resolution as the reference array, SWIR. Fusing all data on to the same grid allows for any follow-on statistics or machine learning model to directly interpret each voxel of training data discretely.The 2D SEM images are roughly the same resolution and have the same pixel dimensions so the data fusion process is insignificant in what it changes. The 3D arrays, however, are quite different with the XCT resolution at 29.96 x 29.96 x2 9.96 μm before fusion and 125 x 125 x 50 μm after the fusion process.

The 3D arrays, however, are quite different with the XCT resolution at 29.96 x 29.96 x2 9.96 μm before fusion and 125 x 125 x 50 μm after the fusion process. Figure 20 below shows the difference in resolution of an XCT slice before (a) and after (b) the fusion process with the same layer in the SWIR array (c).Figure 20: XCT before (a) and after (b) fusion with the same layer in the SWIR array (c).6.0 Discussion of Results and ConclusionsThis study outlined a registration technique that can be used to rapidly register and fuse datasets between multiple sensors and input types and measure the registration accuracy. The process works for both 2D and 3D datasets from multiple data modes and can be extended to other sensor modalities including on- and off-axis optical and thermal images.

The process works for both 2D and 3D datasets from multiple data modes and can be extended to other sensor modalities including on- and off-axis optical and thermal images. The publicly available platform, DREAM.3D, provides repeatability and scalability to this technique so the process can be used for analytics in future work.A point-to-point method has been outlined based on the error distance between matched corresponding points in two-point clouds, and a second point-to-plane method has been outlined based on the error distance between points in one-point clouds to the surface of target dataset. The error metric can be used in DREAM.3D in any type of 2D or 3D dataset, making it easy to objectively judge registration performance in the same units as the dataset.For a point-to-point error metric to be effective, it is vital to select an extremely high number of points in the target point cloud for an effective point-to-point error metric.

The point- to-point error metric converges as the number of points in the target are increased but this study was never able to get enough points in a point cloud to reach the MAE of the point-to-plane method. As more points are added, computing time is exponentially increased. The point-to-plane error metric eliminates this concern because the target is a triangle surface and there is no sampling required.1449Figure 21 provides a visualization of the errors at 10,000 randomly chosen SWIR points using point-to-plane method at the final registered position. The values at each point reflect the point-to-plane error for each point in the SWIR array from the XCT surface with blue and red both reflecting larger error, while white reflects small error (error units are in μm). The largest error is seen at the bottom of the part because the because the SWIR camera did not record during the first ten layers of the build.

The largest error is seen at the bottom of the part because the because the SWIR camera did not record during the first ten layers of the build. This visualization does indicate that the IR registration has a very slight shift along the negative-Y and positive-Z axis. This may be due to the misalignment of arrays in the first ten layers or the inability to apply a calculated transformation matrix to the image geometries. Future development to apply calculated transformation matrices to image geometries is expected to improve registration accuracy.Figure 21: Visual representation of the error in the final 3D registration.There is always some variation in the MAE because the down sampled point clouds are randomly chosen. Each time a new down sampled point cloud is selected there will be a slight difference in the MAE. This variation increases as the number of points in the point clouds decrease because the impact that each randomly chosen point has on the whole result is increased.

This variation increases as the number of points in the point clouds decrease because the impact that each randomly chosen point has on the whole result is increased. This is particularly noticeable in the figures where the down sampling is 0.001 (roughly 23 points) for SEM images or 10 points for the 3D images.The registration accuracy for all eight final SEM positions is 2.4 μm, found with the point- to-plane error metric. The length and width of each pixel in the images are 1.25 μm meaning that the SEM images were registered within two pixels. Future work will be done to combine the eight SEM imaging positions for an improved defect labeling technique. Registration accuracy is vitally important to this work.The registration accuracy for the 3D registration was 75 μm. The voxel size of the fused 3D dataset is 125x125x50 μm so the registration is within a voxel. Future work will be done to train a machine learning model on the fused data using SWIR parameters as model.

Future work will be done to train a machine learning model on the fused data using SWIR parameters as model. A method to label defects in the XCT array is needed for ground truth labels in supervised machine learning. Registration accuracy is vitally important to this work as well.DREAM.3D currently cannot apply the transformations that are computed internally. Further development of the Apply Transformation to Geometry filter to work with image-type datasets and apply interpolation would fully integrate registration into a single application and the process would be faster. Adding this capability to DREAM.3D should increase registration accuracy because there is known variation between the applied transformation and the calculated optimal transformation.1450The 3D images were directly aligned to each other and the 2D images were aligned to a common target.

Using this registration technique and error metric future, work can be done to compare the direct alignment of 3D images to aligning to a common target such as a CAD profile.7.0 AcknowledgementsThis work was funded by the Center for Aerospace Manufacturing Technologies and The Boeing Company. The author is particularly grateful to Sean Donegan, at the Air Force Research Lab, for providing guidance on DREAM.3D and Jeremey Eagan, at The Boeing Company, for general computer science support.8.0 References[1] S. Klein, M. Staring, K. Murphy, M.A. Viergever, J.P. Pluim, Elastix: a toolbox for intensity-based medical image registration, IEEE Trans. Med. Imaging 2009, 29, 196–205.[2]J.B. Forien, N. Calta, P. DePond, G. Guss, T. Roehling, M.

AGE TRANSFORMATIONS AND PRINTING OF PLASTER LAYERS IN SPIRAL GROWTH MANUFACTURINGC. Hauser+, D.M. Lewis†, M. Dunschen*, M. Egan+ , C. Sutcliffe+. +MSERC, Department of Engineering, University of Liverpool, UK †Department of Colour and Polymer Chemistry, University of Leeds, UK *FreeSteel, Liverpool, UK (www.freesteel.co.uk)AbstractSpiral growth manufacturing (SGM) is a high speed rapid manufacturing technique in which objects are built up, layer by layer, by simultaneously depositing, levelling and selectively consolidating thin powder layers onto a rotating build platform. The size and position of the jetted droplets are mapped by the position and greyscale level of pixels within an 8 bit greyscale bitmap image. This paper reports on the development of software in which mathematical algorithms apply geometric transformations to images in preparation for printing onto a rotating substrate.

This paper reports on the development of software in which mathematical algorithms apply geometric transformations to images in preparation for printing onto a rotating substrate. In support of this work, dimensional accuracy measurements of printed images and methods to correct radial print density variations are reported. The accuracy of printed images were found to be within ±0.2mm of their predicted size. The experimental work is briefly extended to the direct printing of plaster layers, formed by mixing two reactive ink solutions.IntroductionThree Dimensional Printing (3DP) was developed at the Massachusetts Institute of Technology in the late 1980’s [1-3]. It is a layered manufacturing process that builds parts by the repetitive deposition of thin horizontal powder layers, regions of which are selectively solidified by inkjet printing a binder from a scanning print head. The application of the binder also bonds consecutive layers together.

The application of the binder also bonds consecutive layers together. The build process is completed with the removal of unbound supporting powder followed by part infiltration (e.g. resin or wax) to increase durability [4]. As with most layer manufacturing technologies, the build methodology is fundamentally a start-stop process since the deposition and processing of each layer occurs sequentially. Hence, the build cycle can be slow, particularly if the build area is large or hardware needs positioning i.e. print heads and powder feeders. Therefore, this technology has largely remained a design prototyping tool or for the manufacture of short run components and tooling.

print heads and powder feeders. Therefore, this technology has largely remained a design prototyping tool or for the manufacture of short run components and tooling. Nevertheless, commercial 3DP technologies are extremely fast in comparison with many other layer manufacturing processes and are capable of building parts at up to 2-6 layers per minute [5].In an attempt to decrease build time, improve system efficiency and broaden the range of viable materials, The University of Liverpool (UoLiv), in collaboration with The University of Leeds (UoLds) and Inovink Ltd, Sheffield, are developing a 3DP process which prints layers of reaction hardenable material directly onto a rotating substrate. During the rotation the substrate continuously falls away from arrays of stationary print heads allowing layers to form one on top of the other. Hence, the substrate rotates in a spiralling motion with a pitch equating to the more familiar layer thickness.

Hence, the substrate rotates in a spiralling motion with a pitch equating to the more familiar layer thickness. The print heads used are OmniDot 760 industrial print heads (XaarJet Ltd, Cambridge, UK). They are wide format (53.8mm swathe) and can print a full swathe in a single pass; this eliminates the need to raster them for area coverage. The inks are two-part solutions, jetted from different print heads which combine on the substrate and react to form calcium sulphate (Plaster-of-Paris). These inks are currently subject to patent review and so only their use will be discussed here. By printing material directly it removes the need to depositpowder layers. This gives the ability to create a layer approaching the maximum print speed of the print head (~500mm/s linear print speed or ~30 layers/minute using the SGM configuration discussed in Equipment and Experimental Methods).

For comparison, deposition of plaster powder layers onto a rotating substrate proved difficult above 12rpm and was limited by poor powder flow properties. The results of this work will be the focus of a future paper.To successfully print an image using SGM, images must first undergo a series of geometric transformations to correct for the annular distortion and parallax shift (a consequence of the dual nozzle row design of the Xaar print heads) caused by printing onto a rotating substrate. This paper reports on the development of computational algorithms to perform these transformations. Python programming language was used in the development of the algorithms [6,7]. The dimensional accuracy of transformed images was tested by experimentation in which images of different size, shape, complexity and radial placement were printed onto paper using ink formulations similar in composition to the reactive inks but with dyes added for print clarity (although non reactive in this instance).

A corrective binary FM screening algorithm was also developed and tested which introduced white noise to monochrome images to homogenise radial print density variations. These experiments were then extended to the printing of the reactive inks. At the time of writing this paper, the reactive inks were not compatible with the OmniDot 760 print head. As a consequence, development work was conducted on a HP DeskJet 5550 printer.Image TransformationsA digital image in which pixels are defined in Cartesian space will, when printed onto a rotating substrate, undergo a polar distortion. The effect of this distortion, by its simplest interpretation, converts horizontal lines to circular arcs and vertical lines to rays radiating out from the origin of the rotation (see Figure 1a). In the process, a typical pixel, Pa, in the x,y plane will be distorted into a curvilinear pixel, Pd, represented in the figure in the u,v plane.

In the process, a typical pixel, Pa, in the x,y plane will be distorted into a curvilinear pixel, Pd, represented in the figure in the u,v plane. Therefore, the first computational step is to remap pixels from Cartesian to Polar coordinate positions. The coordinate system used in the calculations starts at, = , in the upper left corner of an image and the coordinates refer to positions between pixels. The position of pixels, Pt, in the transformed image are described in a similar manner but are represented by the Cartesian coordinates, u,v. The position of pixels in the printed transformed image are again described in x,y. The baseline print density in this work is 360 dpi. This is the radial print density, dpix, and is calculated from the nozzle density across the print head (14.2 nozzles/mm).

The baseline print density in this work is 360 dpi. This is the radial print density, dpix, and is calculated from the nozzle density across the print head (14.2 nozzles/mm). Finally, all measurements in this work are in pixels unless otherwise stated.x=0,y0To start, an 8 bit greyscale or monochrome image is numerically positioned so that its central horizontal axis aligns with an arbitrary radial line and the first column of pixels in the image are at a distance, cx + px. Where, cx, is the radial distance of the first nozzle in the print head and ,px, is the position of the image within the print swathe. In the following calculations, px = 0. The size of the image after the polar transformation is then calculated.

In the following calculations, px = 0. The size of the image after the polar transformation is then calculated. In the image size calculation, the width, su, of the transformed image is a measure of the distance between the pixel closest to and furthest away from the centre of rotation and is given by (see Figure 1b):su=r a np =−cxwhere,a npr==(cx+sx−21 )+((sy/ 2))2 −1(1)(a)(b)Figure 1: (a) The affect of a polar distortion on a pixelated image and (b) calculating the width, su, of a transformed image.The value, -1, is included in equations 1 because distances are taken from between pixels i.e. the coordinates of the first pixel in the image are (0,0).

the coordinates of the first pixel in the image are (0,0). The height, sv, of the transformed image is dependant on the radial position, cx, of the image to be printed and is the product of the length of the arc, vp, subtended by the original image height, sy, during the rotation (see Figure 2a) and the ratio of the circumferential print density, dpiy, at, cx, and the radial print density, dpix:sv=dpidpiyx⋅vp(2)where,pvcxα= ⋅andα⎛ 2 sin = ⎜ ⎝−1sy cx 2⎞ ⎟ ⎠(3)In the above equations, α, is measured in radians, dpix, has a fixed value of 360 dots per inch (see Section 2.2) and, dpiy, is interpolated from values measured experimentally (see Section 4). If max, is the width of the available print swathe, then an 8 bit greyscale transformation image of size, su, sv, is created and all cells are initialised with a white pixel. Each pixel, Pa, from the original image is then mapped into a cell in the transformation image.

Each pixel, Pa, from the original image is then mapped into a cell in the transformation image. The mapping is governed by a polar transformation algorithm and uses the radial distance, a, and the subtended angle, aθ , to calculate the position coordinates,0 su ≤≤S, where, Smaxapr, extending from the centre of rotation to each pixel, P tup,tvp, of each mapped pixel,t aP(see Figure 2b).()(a)(b)Figure 2: (a) Calculating the height, sv, of a transformed image and (b) Mapping of pixel positions in polar transformation calculations.The mapping of pixels from one image to another has been accomplished in two different ways; a forward polar transformation and a reverse polar transformation. The reason for the two approaches, including the differences in both methods, will become clear in the following paragraphs.

The reason for the two approaches, including the differences in both methods, will become clear in the following paragraphs. In the forward polar transformation operation, the calculation starts at pixel, P(a=1)(0,0), in the original image and marches through each pixel in number order (from left to apr , and the subtended angle, θa, from the right and top to bottom).

When printing onto a rotating substrate, the print density, dpiy, at each nozzle changes because the tangential velocity of the substrate at each nozzle is different. Hence, a constant offset, nzroff, across the print head would cause a parallax shift in the printed image (see Figure 4b). The parallax angle, δt, when printing a square of size, su, sv, (see Equations 1,2) positioned in the print swathe at, px, is given by;δ t=tan− ⎛ 1 ⎜ ⎝2(cxd +nzr px+su)⎞ ⎟ ⎠Since the OEM software does not support a variable offset across the print head, each column of pixels, Puc, in the transformed image must be remapped using vertical displacements in, ±v, direction to achieve the required delay. For alignment purposes, the remapping algorithm is geared to implement a delay in one row of nozzles and an advance in the other so that the printed image axis, sx, is parallel to the centre line of the print head.

Puc=0, Puc=2, Puc=4...Puc=762, and all columns of pixels printed by the nozzles in, nzr2, move in the, +v, i.e. Puc=1, Puc=3, Puc=5…Puc=763. The directions can be reversed if the direction of rotation of the substrate and/or print head setup is changed. The nozzles are also staggered, in a repeating pattern of three, across each nozzle row (see Figure 4a). However, the nozzle stagger is significantly less than the nozzle pitch; the latter being used to determine the baseline print density, dpix. When printing at, dpiy = dpix = 360, the required offset to compensate for nozzle stagger would equate to 1 6 pixel (0.0235mm /0.141mm); this would be impossible to achieve since only whole pixels can be moved in the transformation software. Full compensation would first be achievable when printing at, dpiy=2160 (360x6).

Full compensation would first be achievable when printing at, dpiy=2160 (360x6). Since, the effect of nozzle stagger on printed image accuracy is extremely small across relatively large changes in, dpiy, steps required to compensate for it are not considered in this work.Figure 5: Calculating maximum and minimum nozzle row offset values, coff1, and coff2.Image ScreeningExperimental results (see Figures 6/9) illustrated that the print density, dpiy, at eachnozzle reduced with increasing radius according to the tangent function;tany θ= xSince the length, y, is an arbitrary constant, x=cx+ px, and using a suitable scaling factor, X, chosen by curve fitting experimental and calculated data, the print density, dpiy ,across the available print swathe can be evaluated using;(10)ydpi=1 X cx− ⋅Visually, this effect caused a monochrome image to be printed with shades of grey varying from dark shades at, cx, to lighter shades at, (cx + px + su).

To offset radial changes in print density two algorithms were developed. The first algorithm adjusted the grey level intensity of pixels across the image to offset the measured changed in intensity. The Xaar OmniDot 760 print head can produce 5 different drop sizes which directly relate to a small number of discrete pixel intensities distributed across all levels of grey. However, for pixel intensities which fall between these values the associated print head software applied a print pattern to give a perceived impression of grey. The second algorithm made use of digital halftoning techniques to apply FM screening and remove pixels (switch to white) from an image. To start, each column of pixels in an image was grouped into cells. The number of pixels in each cell to be switched was then calculated according to equation 11. The calculated number of pixels were then systematically removed using an equal spacing. This created an image with strong patterning (see Figure 6(i)).

The calculated number of pixels were then systematically removed using an equal spacing. This created an image with strong patterning (see Figure 6(i)). To removed patterning, the next step was to randomise the final position of each switched pixel by displacing it within a given tolerance in ±x and ±y. Figure 6(ii), shows pixel placement with a tolerance of ±(x = 15) and Figure 6(iii) shows pixel placement with a tolerance of ±(x = 15) and ±(y = 10).iiiiiiFigure 6: Graph showing experimental and calculated values of changes in print density, dpiy, with radial position, cx. Included are images of a 600x600 pixel square polar transformed then FM screened (i) systematically, (ii) with noise of amplitude ±(x = 15pixels) and (iii) with noise of amplitude ±(x = 15pixels) and ±(y = 10pixels).(11)Equipment and Experimental MethodsAn SGM machine, designed and built at the UoLiv, was used for all experimental work described in this paper (see Figure 7).

The machine consisted of a 500mm x 100mm high rotating build drum, in which, arrays of stationary print heads were positioned to facilitate image printing on the floor or the drum. Figure 7 shows two SGM configurations. Figure 7a shows one array of three print heads in combination with a powder feeder and wiper blade. Figure 7b shows two arrays of three print heads. The latter configuration has been used in this work although both two have been shown to highlight system flexibility and to show a typical configuration for powder deposition. A detailed description of the print heads are given in image transformations section of this paper. Two of the three print heads (see Figure 7b) in each array are positioned at different radii, and both centralised on two arbitrary radial lines. This gives an effective wide format print swathe of approx 1275 pixels (90mm) with an overlap region, ox, of 253 pixels and a starting radius, cx, of 1605 pixels (113mm).

This gives an effective wide format print swathe of approx 1275 pixels (90mm) with an overlap region, ox, of 253 pixels and a starting radius, cx, of 1605 pixels (113mm). The third print head is positioned at the same radii as print head one. This gives the functionality to overprint using the reactive inks. Although not discussed in detail here, the transformation and screening software described earlier is part of a software suite that also includes motion control software and software to position and prepare images for printing across multiple print heads.During each rotation the drum continuously falls away from the print heads. The drop is driven by a ball screw. The rotational speed of the ball screw is synchronised with the rotation of the drum to give a continuous spiralled layer with a constant pitch. For evaluation purposes, the term ‘layers’ is still used in this work; where one layer is defined by the completion of one complete revolution of the build drum.

For evaluation purposes, the term ‘layers’ is still used in this work; where one layer is defined by the completion of one complete revolution of the build drum. Hence, the pitch is controllable and its magnitude gives a fixed point of reference between the top of the build and the print head nozzles. A rotary and linear encoder attached to the drum and ball screw respectively monitors positional changes. The rotary encoder clock pulse is also used to synchronise events in the print heads.Monochrome bitmap images, included screened images, of squares, rectangles and circle that varied in size were transformed and printed at different values of, px, to test for printed size accuracy and reproducibility. The rotational speed ranged from 1 to 35 rpm. All images were printed onto strips of paper that were taped to the floor of the build drum. The ink used in the experiments was an aqueous based ink with blue dye added for print clarity. This ink was a precursor to the reactive inks.

The ink used in the experiments was an aqueous based ink with blue dye added for print clarity. This ink was a precursor to the reactive inks. The size and diagonal deviation (squareness) of printed images were measured using a microscope with a Vernier scale attached to the stage. The results were then compared to the predicted image size. The predicted size of an image is calculated using the baseline print density of 360 dpi (see image transformations section). Greyscale shading of printed images was approximated by digitising the printed image (using an Epson 2480 photo scanner) then measuring the grey level of pixels at equal intervals across the image using Adobe Photoshop™ software (using the eyedropper tool set to a 5x5 average pixel sample).Multiple layer passes using the reactive inks was carried out using a HP DeskJet 5550 printer (17pl drop size @ 1200dpi) [8]. The reactive inks were printed onto paper, one pass for each ink, up to a maximum of 4x2 layers.

The reactive inks were printed onto paper, one pass for each ink, up to a maximum of 4x2 layers. Attempts to print more layers was hindered by the raised image passing through the paper feed mechanism of the printer. The two inks were contained in different cartridges and were interchanged after each print pass.Figure 7: Schematic diagram showing SGM build drum, position of print head arrays and powder processing equipment.Results and DiscussionSoftware tools have been developed to transform images in preparation for accurate reproduction onto a rotating substrate. A Typical example of a transformed image and its printed form is shown in Figure 8. In Figure 8, the print quality from print head 2 is deliberately low to highlight the circular path taken by the substrate as it passed under the print head. There are still some issues when trying to seamlessly rejoin an image when printed across two print heads.

There are still some issues when trying to seamlessly rejoin an image when printed across two print heads. However, the authors believe this is principally a print head alignment issue rather than being associated with software errors. The size accuracy in, x, of transformed images printed onto paper was ±0.1mm of their predicted size. The printed image size accuracy in, y, is less accurately defined because it has to be scaled according to equation 2 and is a function of radial position, image size and encoder divide value, end. Dimensional measurement analysis on a wide sample of regular quadrilaterals printed at different radial positions were typically within ±0.2mm (this value appeared to be independent of image size). The diagonal deviation (squareness) of printed images was typically ±0.5%.The screening algorithm was used to model the radial print density variation across the total print swathe, pxmax.

The diagonal deviation (squareness) of printed images was typically ±0.5%.The screening algorithm was used to model the radial print density variation across the total print swathe, pxmax. For the print swathe settings described in this work, the total print density drop was calculated to be 35% (see Figure 9) and the rate of change with, px, followed the curve shown in Figure 6. For comparison, the grey level intensity drop measured using Photoshop™ was 27% for the same sample image. The disparity may result from the level of software interpretation when digitising printed images. Visually, screened images ordinarily looked to be corrected, although perhaps marginally over corrected at lower values of, cx (see Figure 9b). In an attempt to quantify the efficacy of the screen algorithm, screened printed images were digitised and examined using Photoshop™.

In an attempt to quantify the efficacy of the screen algorithm, screened printed images were digitised and examined using Photoshop™. Results suggested the print density variation was reduced to 4% across the total print swathe.Figure 8: (a) Digitised 8 bit greyscale image, (b) transformed image and (c) image printed using SGM apparatus.Figure 9: SGM printed image of size 1200x700 px (a) without screen and (b) with screening.Two part reactive inks were demonstrated by overprinting the same image twice (see Figure 10); each pass using a different ink. The reaction was instantaneous giving solid precipitates of plaster which fully dried to give a relative smooth solid layer. The average layer thickness was 25μm and the density of the plaster image approached 70% (compared with compacted plaster powder); density measurements were taken using four layer samples (4x2 passes). For comparison, current 3D printing technologies tend to achieve 50-60% density from the machine.

For comparison, current 3D printing technologies tend to achieve 50-60% density from the machine. The printed plaster was also unexpectedly flexible and showed no signs of cracking when bending the paper; even to a sharp right angle. Trials are currently in development to adapt these inks for use in forthcoming versions of the Xaar print heads. It is anticipated that the SGM system will give greater flexibility over the choice of layer thickness and material density since the drop volume and print density is more adaptable than in the HP printer.ConclusionsThis paper has briefly reviewed developments in three dimensional printing and highlighted the need to develop the technology for more efficient and faster layer construction. To this end, the development of the spiral growth manufacturing process for the three dimensional printing was conceived. However, printing onto a rotating substrate poses a number of issues; namely the accurate reproduction of digital images.

However, printing onto a rotating substrate poses a number of issues; namely the accurate reproduction of digital images. Therefore, the work of this paper focussed on a number of transformation and screening algorithms to prepare 8 bit greyscale and monochrome images for rotational printing using arrays of wide format, dual nozzle row, industrial print heads. Two transformation algorithms have been presented, one to polar transform the digital image and one to offset columns of pixels for printing by nozzles in two(a)(b)Figure 10: Reactive ink printing (a) single layer image and (b) multiple layer image (4 passes).different parallel rows. Image reproducibility, accuracy and squareness were within reasonable limits for printing reactive inks, or binders onto powder (accuracy ±200µm, squareness ±0.5%). It is currently difficult to highlight areas where accuracy improvements can be made in the transformation algorithms.

It is currently difficult to highlight areas where accuracy improvements can be made in the transformation algorithms. However, it is expected that more precise mechanical alignment of the print heads may lead to tolerances that equate similar to images printed on a linear conveyor. Results from the screening trials were inconclusive. Visually, the radial greyscale gradient, which is clearly visible on all images printed on the SGM apparatus appeared to be corrected after images were screened. However, the impact of radial print density variations, if any, on build quality cannot be fully evaluated until such time when reactive inks can be printed using SGM, and build geometries mechanically tested. The reactive inks show great promise, particularly for use in SGM, and would exploit the speed potential of a rotary system (minimum of 30 layers per minute build rate).

The reactive inks show great promise, particularly for use in SGM, and would exploit the speed potential of a rotary system (minimum of 30 layers per minute build rate). Preliminary tests on the printed plaster layers suggest the material has good strength, density, wear resistance and bonds extremely well to underlying layers.AcknowledgementsThis work has been supported by grants from the UK Engineering and Physical Sciences Research Council (EPSRC). In particular, thanks must be given to Martin Dunschen and Julian Todd from FreeSteel, UK, for their help in developing the transformation algorithms. One of us (C.H.) is grateful to Martin for his training in computer programming.References[1] [2][3][4][5][6] [7] [8]Jacobs, PF (1996). Stereolithography and other RP&M technologies. ASME Press, New York. Sachs, E., Cornie, J., Brancazio, D., Bredt, J., Curodeau, A., Fan, T., Lauder, A., Lee, J. and Michaels, S. (1993).

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceIMMISCIBLE-INTERFACE ASSISTED DIRECT METAL DRAWINGLi He1,2, Fan Fei1,2, Wenbo Wang1,2, Xuan Song1,2*1Department of Industrial and Systems Engineering, University of Iowa, Iowa City, IA52242, USA2 Center for Computer Aided Design, University of Iowa, Iowa City, IA 52242, USAAbstractState-of-the-art metal additive manufacturing (AM) techniques construct a three- dimensional (3D) structure through sintering or melting dry metal powders in a layer-by-layer fashion, which consequently results in some typical manufacturing defects in the final structure, such as residual stress and highly-orientated microstructures.

To overcome these defects, we present a new low-cost metal AM process, named Immiscible-interface assisted Direct Metal Drawing (II-DMD), which fabricates self-supported and isotropic 3D metal structures by continuously extruding a metal colloidal suspension within a second immiscible matrix colloidal suspension. The shape of the metal colloidal suspension is stabilized due to the presence of an immiscible interface between the two colloidal suspension systems. Dense metal structures can be achieved via post-consolidation of the self-stabilized metal-matrix systems, including liquid-phase drying and metal-phase sintering. In this article, the immiscible-interface-assisted self-stabilization mechanism is studied. The post-consolidation processes are discussed. Several test cases were fabricated and characterized.Keywords: immiscible interface, direct metal drawing, colloidal suspension, self-stabilization, heat treatment1.

Several test cases were fabricated and characterized.Keywords: immiscible interface, direct metal drawing, colloidal suspension, self-stabilization, heat treatment1. IntroductionTraditional metal additive manufacturing (AM) processes [1] build a three-dimensional (3D) metallic structure through selectively sintering or melting dry metal powders in a layer-by- layer manner using a high-energy beam. The layerwise fabrication method used in those processes results in many typical manufacturing defects in final structures, such as the staircase effect, residual stress and highly-orientated microstructures. Specifically, the staircases can reduce the surface finish and introduce stress concentration; residual stress and highly-orientated microstructures caused by the localized high-temperature heating [2, 3] reduce mechanical properties of final objects.

In addition to these drawbacks, fabrication costs of existing metal AM processes are substantially high, due to required high-energy sources, which limit broader application in industry.In this paper, we report a new AM process for fabricating metallic structures ( in particular lattice structures), i.e., immiscible-interface assisted direct metal drawing (II-DMD), which offers the potential to mitigate the inherent defects associated with layerwise fabrication approaches. In this new process, a metal-powder-based suspension is first deposited continuously in 3D spaceCorresponding author.2184within a secondary matrix suspension and then uniformly densified through heat treatment. By tailoring the two colloidal systems, a suitable immiscible interface can be established between the metal-powder suspension and the secondary matrix suspension, which consequently enables the self-stabilization of the deposited 3D geometry.

Similar fabrication techniques [4-12] have been used to create microvascular networks[12], soft sensors[10], robots[11], and other 3D structures, which are however still limited to polymer materials (e.g. hydrogel) and have not yet been attempted to build metal structures. The paper is organized as follows. Section 2 describes starting materials used in the II-DMD process; section 3 introduces a fabrication system; section 4 discusses a mathematic model for the nozzle extrusion of the II-DMD process; section 5 demonstrates several test cases and their properties; conclusions and future work are discussed in section 6.2. Starting materialsIn this research, bronze powders are used as a model material to study the II-DMD process. Aluminum oxide is selected to form a secondary matrix suspension due to its high sintering temperature (> 1300oC).

Aluminum oxide is selected to form a secondary matrix suspension due to its high sintering temperature (> 1300oC). Bronze-powder-based suspensions are prepared by mixing as-received bronze powders (10% Tin, 90% Copper and 0.07% Phosphor, Mesh: 325, Chemical Store Inc. Clifton, NJ, USA) with silicone oil at a specific concentration in a magnetic stirrer for ~15 minutes. The resulting bronze-powder suspension needs to exhibit a shear-thinning behavior with yield stress to achieve controllable extrusion through a pressurized dispenser. Moreover, the shear elastic modulus and yield stress of the achieved bronze-powder suspension should be an order of magnitude larger than those of the used secondary matrix suspension to ensure that the suspension maintains its shape after it exits the nozzle [13].

A secondary matrix colloidal suspension is prepared by mixing alumina powder (GE6, Baikowski, Charlotte, USA) with deionized water at a specific concentration in a magnetic stirrer for ~15 minutes. Similarly, the achieved matrix suspension should exhibit a shear-thinning behavior with suitable yield stress to support the extruded bronze-powder suspensions [14]. Additionally, the matrix suspension needs to be thixotropic such that crevices generated by nozzle movement can be immediately closed through a time-dependent backflow.Silicon oil and water are employed as the dispersion medium of metal suspensions and secondary matrix suspensions respectively to introduce an immiscible interface [15, 16] between bronze particles and alumina particles, which is required to achieve the self-stabilization of a deposited 3D shape. More details are discussed in section 3.3. Process DesignA schematic representation of the II-DMD process is shown in Figure 1.

More details are discussed in section 3.3. Process DesignA schematic representation of the II-DMD process is shown in Figure 1. A pressurized dispensing syringe is mounted on a three degrees of freedom (DoF) gantry system. The syringe is loaded with a bronze-powder suspension and is controlled to continuously dispense the suspension within an alumina matrix suspension in a crucible, as depicted in Figure 1a.1. As the bronze- powder suspension is deposited, an immiscible water-oil interface is established between the bronze and alumina particles, which locks the bronze and alumina particles in position and maintain the 3D shape of the deposited bronze-powder suspension, as shown in Figure 1a.2.2185Rheological properties of each suspension have been carefully tailored to avoid deformation of the deposited 3D shape during the fabrication (e.g., fragments, bead up, diffusion, or sedimentation).

After a bronze-powder suspension with the desired geometry is achieved, the crucible is heated in an oven to dry the bronze-powder and alumina-matrix suspensions. The drying schedule is given in Figure 1b. Through this step, a dry bronze-powder compact with the desired geometry is formed, surrounded by a nearly dry alumina powder bed (refer to Figure 1a.3). The alumina powder bed still contains a small amount of silicon oil that diffused from the bronze-powder suspension during drying. After that, both the bronze-powder compact and the alumina powder bed are heated together in an argon furnace. The heating temperature schedule is given in Figure 1c. Since the used temperature (i.e., 950 ºC) is much lower than the sintering temperature of alumina powder (e.g., 1400 ºC), only bronze particles are sintered/melted, with the alumina matrix powder remaining loose. In addition, during the high temperature sintering, the residual silicone oil in the alumina powder bed is also decomposed.

In addition, during the high temperature sintering, the residual silicone oil in the alumina powder bed is also decomposed. Finally, the alumina matrix powder is removed and recycled for future use, leaving the sintered bronze body as the final product.Figure 1. (a) Schematic diagram of II-DMD; (b) temperature schedule for drying; (c)temperature schedule for sintering.4. Process ModelingThe required solid loadings of bronze-powder suspensions and alumina matrix suspensions in the II-DMD process were experimentally determined as 94wt% and 40wt%, respectively, which yield the optimal stability and printability during the process. These materials were used to study the relationship between extrusion pressures, printing speed and extruded filament size (with a fixed nozzle size of 20 Gauge).

These materials were used to study the relationship between extrusion pressures, printing speed and extruded filament size (with a fixed nozzle size of 20 Gauge). Figure 2 shows the experimental results and fitting models.2186As seen in Figure 2a, the achieved metal filament size in an alumina powder bed after drying is proportional to the applied extrusion pressure. In contrast, the achieved metal filament size is inversely proportional to the square root of printing speed, as shown in Figure 2b. Based on the experimental results, bronze filament sizes achieved through a 20 Gauge nozzle can be expressed as the following equation:−0.5In which:   𝑦𝑦 ≈ 250√2 ∗ 𝑝𝑝 ∗ 𝑣𝑣y is the filament size (unit: µm); p is the extrusion pressure (unit: psi); and v is the nozzle printing speed (unit: mm/s).(a) (b) Figure 2. Effects of (a) extrusion pressure and (b) printing speed on extruded filament size.5.

Effects of (a) extrusion pressure and (b) printing speed on extruded filament size.5. Test CasesIn this section, several test cases were fabricated to showcase the presented II-DMD process. Figure 3a shows a bronze lattice structure, which can be used as structural frames of micro robots. Figure 3b shows a stent structure, which can be applied to prop up a pathological changed blood vessel in the human body if shape memory alloy powder is used.To further demonstrate the capability of the II-DMD process, two cone springs with different sizes were printed and shown in Figure 3c. An optical microscope image of the external surface of the springs is shown in Figure 3d. The rough surface finish is mainly caused by a large bronze particle size used in the process. SEM images of the spring specimens are shown in Figure 3e and 3f.

The rough surface finish is mainly caused by a large bronze particle size used in the process. SEM images of the spring specimens are shown in Figure 3e and 3f. It can be seen that small pores with diameters around 10 to 40 µm exist in the cross section, which may be caused by inappropriate selection of bronze particle size and drying parameters. In Figure 3f, the metal cross section is dense at a macroscopic level, which contributes to relatively good mechanical properties of printed parts. According to a tensile test as shown in Figure 3g, bronze specimens fabricated by our method have a yield strength of 106 MPa and an ultimate strength of 148 MPa.2187(1)Figure 3. Test cases: (a) a lattice structure, (b) stent structure, (c) cone springs, (d) opticalmicroscope image of springs, (e) and (f) SEM images of printed samples, (g) tensile test of a fabricated specimen.6.

Conclusion and Future WorkIn this article, we report a new and low-cost metal 3D printing process, in which an immiscible interface and unique rheological behaviors of highly-loaded metal and ceramic colloidal suspensions are utilized to build complex metal lattice structures. It can potentially fabricate numerous metal materials without generating staircase effect and the need of support structures, and the microstructures of achieved metal parts can be homogeneous and isotropic. Bronze materials are used as a model material to demonstrate the process, with alumina powder as a matrix material. In the future, problems such as rough surface finish and porosity still need to be solved.

In the future, problems such as rough surface finish and porosity still need to be solved. Fundamental physics behind the II-DMD process will be studied and more materials will be tested.ReferencesGu, D., Meiners, W., Wissenbach, K., and Poprawe, R., 2012, "Laser additive manufacturing of [1] metallic components: materials, processes and mechanisms," International materials reviews, 57(3), pp. 133-164. [2] Strößner, J., Terock, M., and Glatzel, U., 2015, "Mechanical and Microstructural Investigation of Nickel-Based Superalloy IN718 Manufactured by Selective Laser Melting (SLM)," Advanced Engineering Materials, 17(8), pp. 1099-1105.2188Chen, H., Gu, D., Dai, D., Ma, C., and Xia, M., 2017, "Microstructure and composition homogeneity, [3] tensile property, and underlying thermal physical mechanism of selective laser melting tool steel parts," Materials Science and Engineering: A, 682, pp. 279-289. [4] T.

pact Absorbent Rapid Manufactured Structures (IARMS)J. P. J. Brennan-Craddock, G. A. Bingham*, R. J. M. Hague, R. D. Wildman The Rapid Manufacturing Research Group, Wolfson School of Mechanical and Manufacturing Engineering, Loughborough University, Leicestershire, UKAbstractRapid Manufacturing (RM) is increasingly becoming a viable manufacturing process due to dramatic advantages that it facilitates in the area of design complexity. Through the exploration of the design freedom afforded by RM, this paper introduces the concept and initial research surrounding Impact Absorbent Rapid Manufactured Structures (IARMS), with an application in sports personal protective equipment (PPE). Designs are based on the cellular structure of foams; the inherent advantages of the cellular structure are used as a basis to create IARMS that have the potential to be optimised for a specific impact absorbent response.

The paper provides some initial results from compression testing.IntroductionPPE is used in many sports to reduce the risk of injury to the athlete or player. Conventional PPE can be cumbersome and ill-fitting as the design is limited by their manufacturing processes. There is a need for customisation for the individual athlete to take into account variations in body size and shape; this not only serves to reduce potential injuries but also improve player performance.

Utilising the design freedom and geometric complexity of Rapid Manufacturing (RM) techniques, the project aims to create customised PPE that is not only optimised for the specific impact forces and velocities involved within a particular sport but also the overall weight, flexibility and comfort perception of the PPE garment.The Scuta ProjectThe research documented within this paper was conducted under an EPSRC/IMCRC funded Integrated Project conducted at Loughborough University; Tailored Injury Prevention & Performance Improvement for Protective Sports Garments (known as Scuta) [1].

The overall project aims are the optimisation and customisation of Personal Protective Equipment (PPE) utilising Rapid Manufacturing (RM) techniques for a number of test-case contact and ball sports: Tae kwon-do, Cricket and Association Football.Within Scuta, Work Package 4 (WP4): The Generation of 3D Conformal Data is investigating both the specific design of IARMS and potential modelling strategies that allow the generation of conformal and body-fitting RM PPE to be realised.

The work packages contained within the integrated project are responsible for investigating related issues as follows [1]:Work for this paper was undertaken at Loughborough University, however this author has now taken a position within the School of Art & Design, De Montfort University, Leicestershire, UKWP1: Determining Human-Related Impact Intensity During Contact Sports WP2: Perception & Comfort of Personal Protective Equipment WP3: Advanced Modelling of Personal Protective Equipment for Sports WP4: The Generation of 3D Conformal Data WP5: Materials Analysis WP6: Instrumentation & ValidationThe research documented within this paper concentrates on the initial design and compression testing of IARMS manufactured using the Laser Sintering process. The Computer Aided Design (CAD) work was completed using NX5 and Magics.Application and Aim of IARMS DesignThe Tae kwon-do chest protector (known as a hogu) has been selected as an application to focus research.

Tae kwon-do is a martial art that employs punches and kicks, with kicking accounting for 80% of strikes [2]. Kicks can be broadly characterised as swing or thrust kicks, delivering high-speed impacts to a contestant over the course of a match [2]. The aim for IARMS designs in development is to generate personalised, resilient structures tailored to this environment.Design ResearchWP5: Materials Analysis is responsible for investigating and developing new impact absorbent materials suitable for both additive manufacturing techniques and the actual manufacture of IARMS. However, as this research is being conducted concurrently, an initial constraint for IARMS design is the utilisation of commercially available materials.

However, as this research is being conducted concurrently, an initial constraint for IARMS design is the utilisation of commercially available materials. Given this initial design constraint, the design research has been directed at utilising the geometric freedom afforded by additive manufacture to provide the impact absorbent characteristics required, rather than the actual physical properties of the build material. To aid this design process, existing conventional impact absorbent materials, specifically foams, have been investigated.Structure of FoamA wide range of materials are utilised in engineering applications for impact absorption. In personal protection (e.g.: impact absorbing PPE) and product protection (e.g.: packaging), polymeric foams dominate the market [3]. The favourable impact characteristics of foam are derived mainly from its cellular structure which is classified as either open or closed-cell.

The favourable impact characteristics of foam are derived mainly from its cellular structure which is classified as either open or closed-cell. A closed-cell structure consists of walls and struts, forming isolated cells of gas. An open-cell structure consists solely of struts – during manufacture, the cell walls formed between nucleating bubbles are pierced and shrink into the forming struts. This creates a structure where a fluid (ie: air) can flow freely through it [3]. Examples of open and closed-cell foam are shown in Figure 1.Figure 1: SEM photographs of open-cell (left) and closed-cell (right) structure of foam [3,6]Compressive Behaviour of FoamFoams are particularly good at dissipating the energy and reducing the peak force of an impact due to their low density, elastically deforming structure. The cellular structure also lends itself well to absorbing oblique impacts [3]. Foams have a characteristic three-regime compressive behaviour, shown schematically in Figure 2.

The cellular structure also lends itself well to absorbing oblique impacts [3]. Foams have a characteristic three-regime compressive behaviour, shown schematically in Figure 2. Initially, the foam exhibits a linear elastic region where very little energy is absorbed. A wide plateau region follows, with a densification phase where stress rises steeply [4]. The energy absorbed by a foam, per unit volume is shown by the area under the curve in Figure 2. In regards to flexible foams, the manner in which energy is 'absorbed' in the elastic stages of compression is in potential energy stored by deforming the structure.Linear elasticWide stress plateauDensifications s e r t S e v i s s e r p m o CEnergy absorbed per unit volumeCompressive StrainFigure 2 – Characteristic compressive behaviour of flexible foam (based on a diagram from [5])The 'wide stress plateau’ illustrated in Figure 2 is a characteristic of foams; the reason why they make such good energy absorbers.

It is a stage of elastic, localised collapse where weaknesses in the structure buckle [6]. Further deformations will spread from these localised areas until the foam is completely compressed. The final regime of compression sees significant stiffening of the foam in a phase known as densification [6]. This stage is where plasticdeformation of the foam occurs, reducing its capacity to absorb subsequent impacts. Cells have completely collapsed and the struts of the structure are compressed into contact. The structure now more closely approximates the properties of a solid block.In open-cell foams, energy is absorbed through the bending of the structure’s struts. Closed-cell foams will generally absorb higher impacts as compression of the gas trapped inside cells will contribute to its energy absorption mechanisms [7]. An ideal foam would have a stress plateau that ended at the maximum impact stress associated with the application [3].

An ideal foam would have a stress plateau that ended at the maximum impact stress associated with the application [3]. The foam would not reach the densification phase during normal operating conditions. This would ensure that the product did not suffer fatigue, but it would also ensure that minimal impact energy is transferred to the object that it is protecting.This criterion can be transferred to the design of tailored IARMS. The cellular structure of foam is a suitable starting point for the design of IARMS as it has been shown to be a fine energy absorber. The irregular structure of foam is difficult to replicate in the design of IARMS, so a simplification of the structure has been applied using a model known as the Kelvin structure.The Kelvin StructureThe foaming process generates a three-dimensional cellular structure that is controlled by the principle of minimal surface energy [8].

Kelvin proposed a body-centred cubic lattice of a 14- faced polyhedral cell as a representation of an idealised foam structure [9]. The Kelvin structure is based on a modified truncated octahedron (a polyhedron composed of six square and eight hexagonal faces): the hexagonal faces have zero mean curvature, while the square faces are flat with outwardly curved edges, as illustrated in Figure 3.Figure 3 - A Kelvin cell (with curved faces) and the planar-faced truncated octahedron it is based onThe Kelvin structure has since been used as the basis for finite element analyses (FEA) of foam structures to better understand the relationship between structure and properties [10, 11, 12].

Although originally proposed as a model of the minimal surface energies characteristic of foams, the Kelvin structure has been shown to be a good model of its mechanical properties [10, 11, 12] and thus shows promise as the base for the design of IARMS.Initial IARMS DesignThe Kelvin structure is comprised of large, compressible polyhedra. Due the open, low relative density of the structure, it can be compressed to high percentage strains before structural elements come into contact. An open-cell structure has been adopted to facilitate the removal ofun-sintered powder from samples created. The wide angles between meeting struts means the ‘node’ between them needs only be the same diameter as the struts. Minimising the node allows the structure increased compressibility; large nodes in structures like this are generally solid and reduce the structure’s compressibility.

Minimising the node allows the structure increased compressibility; large nodes in structures like this are generally solid and reduce the structure’s compressibility. Using the Kelvin structure as a base, different geometries can be applied to it to generate a range of designs; one such design has been named the 'straight s trut' design, as illustrated in Figure 4.StrutNodeFigure 4 - 'Straight strut' IARMS designA primary purpose of this design is the minimisation of file size. Minimal polygons required to represent struts to facilitate the generation of physically large structures. Each strut is essentially a triangular prism and is represented by twelve polygons in an .stl file, as illustrated in required to represent a Figure 5. This compares very favourably to the number of polygons ylinder (also illustrated in Figure 5) and significantly reduces file size.

This compares very favourably to the number of polygons ylinder (also illustrated in Figure 5) and significantly reduces file size. cTriangular polygonFigure 5 – Polygon representation of strut geometriesEach node of the structure has fillets applied to it to reduce the chance of crack formation, although the fillets can be removed. The addition of a set of fillets at each node does significantly increase the polygon count of the structure, although compared to that of a strut of circular cross-section the polygon count is still favourable.Crack propagation site at node without filletFilleted nodeFigure 6 – Fillets applied to a nodeBy applying a more complex helix geometry to the Kelvin structure, a more flexible IARMS design is created. The 'helical strut' design is illustrated in Figure 7. The higher geometric complexity in comparison to the straight strut design significantly increases the .stl file size.

The 'helical strut' design is illustrated in Figure 7. The higher geometric complexity in comparison to the straight strut design significantly increases the .stl file size. The profile of the helix that comprises the strut follows a law curve, as shown in Figure 7. This allows multiple struts to meet at a node, ensuring contact only at the node.Law curveLaw helixStrutFigure 7 - 'Helical strut' IARMS designCompression Testing of IARMSInitial compression tests have been conducted on the two variants of the IARMS design, summarised in Table 1. Each variant was manufactured in a blend of 50% virgin/50% recycled PA 2200 material from EOS on a Formiga 100P, although for comparison, a set of straight strut samples were also manufactured in Duraform EX on a Vanguard with HiQ upgrade.

Due to the unknown properties of these newly developed structures, low-rate compression tests were conducted on samples, rather than high-rate impact tests more representative of Tae kwon-do strikes.Design VariantStraight strutStraight strutHelical strutMachine/ MaterialVanguard with HiQ upgrade / Duraform EXFormiga 100P / PA 2200Formiga 100P / PA 2200No. of Samples Manufactured555.stl File Size0.88 MB0.88 MB26.50 MBSample Dimensions75 x 75 x 30mm75 x 75 x 30mm75 x 75 x 30mmCompressive Extension25mm25mm25mmStrain Rate100mm/min100mm/min100mm/minTable 1 – Compression samplesMethodologyThe IARMS samples were compressed on an Instron 3366 Dual Column Testing System at a strain rate of 100mm/min. Tests were controlled by extension and the load measured because the compressive strength of the structures were unknown.

Tests were controlled by extension and the load measured because the compressive strength of the structures were unknown. The test setup is shown in Figure 8 and summarised in Table 1.Compression platenSampleFigure 8 – Flat sample compression test setupEach 30mm thick sample was compressed by 25mm once and the reaction force measured. From this, the compressive stress and compressive strain was calculated and plotted.ResultsCompressive Behaviour of the Straight Strut DesignsThe three regimes of compression that are commonly associated with foam have been observed in the compression of the straight strut samples. As shown in Figure 9, the Duraform EX straight strut samples exhibit an initial, linear elastic response until approximately 15% strain when the structure softens. The structures deform significantly across a wide stress plateau with relatively minimal exertion.

The structures deform significantly across a wide stress plateau with relatively minimal exertion. Finally, at around 75% strain the samples stiffen considerably as the struts that compose the structure come into contact during densification. This is where the peak stresses of these samples occur, between 25 and 30kPa.Figure 9 – Compressive response of straight strut samples (Duraform EX)The PA 2200 straight strut samples exhibit a very similar compressive behaviour, shown in Figure 10. There is a greater variation between PA 2200 samples, however they perform similarly to the Duraform EX samples. Both sets of samples plateau at around 10 to 15% strain, with a compressive stress around 15kPa. Both sets of samples increase in stiffness as they enter the densification stage between 70 and 75% strain.

Both sets of samples increase in stiffness as they enter the densification stage between 70 and 75% strain. The range of peak stresses of the PA 2200 samples is slightly lower than that of the Duraform EX samples, with the exception of one anomalous sample shown with a dashed curve in Figure 10.Figure 10 – Compressive response of straight strut samples (PA 2200)It is during the densification stage that plastic deformation of the samples’ structure begins to occur. The samples do not return to their original height, as shown in Figure 11. Plastic hinges have formed at the nodes of the structure, creating a largely regular deformation pattern throughout it. Figure 11 clearly shows that the plastic deformation has occurred at the nodes rather than in the struts, which remain relatively straight as the structure compresses. Some fracturing occurs around the nodes within the structure; struts have failed where the fillets on the nodes finish.

Some fracturing occurs around the nodes within the structure; struts have failed where the fillets on the nodes finish. The level of actual fracturing within the structure is minimal and is not a significant cause of the structure’s loss of resilience.Figure 11 – Compression of a straight strut sampleCompressive Behaviour of the Helical Strut DesignA different compressive behaviour is observed of the helical strut sample. The helical structure exhibits potentially linear stiffness, replacing the first two regimes of compressionfound with the straight strut samples, with a densification stage beginning at approximately 70% strain.

This is shown in Figure 12.Linear elastic regimeDensification) a P k (350s s e r t S e v i s s e r p m o C300250200Helical strut sample15010050Helix strut samplesStraight strut sample001020304050607080Compressive Strain Compressive Strain (%)Figure 12 – Compressive response of helical strut sample, with the straight strut response shown for comparisonDensification in the helical strut samples occurs earlier than in the straight strut samples because the helical geometry of the struts make contact earlier than the straight struts.Figure 13- Compression of a helical strut sampleThis particular helical strut design absorbs significantly more energy than the straight strut design, an example shown in Figure 12 for comparison. The helical strut design did not recover to full height; as with the straight strut samples some plastic deformation occurred.

The helical strut design did not recover to full height; as with the straight strut samples some plastic deformation occurred. It is less obvious where plastic hinges have formed throughout the helical strut samples due to the complexity of the structures.Conclusions and Further WorkThe research documented in this paper shows through physical testing that the Kelvin structure does approximate the compressive behaviour of flexible foams. The structure designs tested exhibit very different properties. The straight strut samples exhibit a compressivetwo90response similar to that of conventional foams, with the characteristic three regimes of compression. By inserting a different strut geometry into the structure, a completely different compressive behaviour is observed; the helical strut sample displays a two regime compression; a linear elastic phase followed by densification.

The helical strut sample absorbs significantly more energy than the straight strut design.A major limitation in the generation of lattice structures (such as IARMS) for Rapid Manufacture are the file sizes required to represent them. Although the helical strut design is more resilient than the straight strut design, its .stl file is thirty times larger. It is currently unfeasible to generate a structure large enough to form a Tae kwon-do hogu or similar PPE of the helical strut design with conventional CAD. Additionally, the designs discussed in this paper are not capable of withstanding the forces associated with Tae kwon-do at this point.

Additionally, the designs discussed in this paper are not capable of withstanding the forces associated with Tae kwon-do at this point. Designing a strut geometry with properties exceeding that of the helical strut design but with a minimised .stl file size is a key focus of this ongoing research.The samples tested in this study have not been optimised for any particular environment – altering the strut parameters of the designs will effect the structure’s compressive behaviour. A parametric study will be conducted to identify what these effects are and the design of other strut geometries is ongoing. The results suggest that it should be possible to tailor IARMS to specific applications, through the selection of optimised parameters. Additionally, strut parameters or designs could be varied throughout a structure to create functionally graded garments. Although they underwent some plastic deformation, i.e.

Additionally, strut parameters or designs could be varied throughout a structure to create functionally graded garments. Although they underwent some plastic deformation, i.e. they did not reform back to their original shape after compression, the samples did not break into pieces. With optimisation, these structures could potentially be much more resilient.The loading rates applied to the samples in these compression tests are not representative of the impacts associated with Tae kwon-do. Further tests must be conducted in more representative conditions to determine the effects of strain rate on the structures’ compressive behaviour. To generate conformal IARMS, a means to create body-fitting garments is being investigated.

To generate conformal IARMS, a means to create body-fitting garments is being investigated. Although not covered in this paper, a key aspect of this research is the development of a modelling methodology to generate conformal samples of the structure that will fit to the curves of an individual’s body.References[1] Hague, R., (principal investigator), 2006, EPSRC/IMCRC integrated project number 185, Injury Prevention & Performance Improvement for Protective Sports Garments (SCUTA), start date October 2006 Serina E.R., Lieu, D.K. Thoracic injury potential of basic competition taekwondo kicks. Journal of Biomechanics. 1991;24(10):951-60.[2][3] Gibson L.J., Ashby, M.F. Cellular Solids. 2 ed. Cambridge: Cambridge University Press1997.[4] Wang Z., Ma, H., Zhao, L., Yang, G. Studies on the dynamic compressive properties of open-cell aluminum alloy foams. Key Engineering Materials. 2006;306:905-10.[5] . packaging and energy absorption. Journal of Materials Science.

pact and Influence Factors of Additive Manufacturing on Product Lifecycle CostsC. Lindemann*, U. Jahnke*, M. Moi*, R. Koch*Direct Manufacturing Research Center (DMRC) and Chair of Computer Application and Integration in Design and Planning (C.I.K.), The University of Paderborn, Mersinweg 3, 33098 Paderborn, GermanyAbstractAt first sight the direct costs of Additive Manufacturing (AM) seem too high in comparison to traditional manufacturing. Considering the whole lifecycle costs of parts changes the point of view. Due to the modification of the new production process and new supply chains during a parts lifecycle, producing companies can strongly benefit from AM. Therefore, a costing model for assessing lifecycle costs with regard to specific applications and branches has been developed. The costing model represents the advantages of AM monetary.

Therefore, a costing model for assessing lifecycle costs with regard to specific applications and branches has been developed. The costing model represents the advantages of AM monetary. For the evaluation of this model and the influence factors, different case studies have been performed including different approaches in part redesign. Deeper research is and will be carried out with respect to the AM building rates and the comparability of various AM machines, as these facts are hardly comparable for end users. This paper will present the methodology as well as the results of the case studies conducted over the whole product lifecycle.IntroductionOne of the critical success factors for the additive manufacturing technologies can be seen as the costs for the manufacturing process compared to traditional manufacturing [EKW+12]. Especially those potential users / customers that are not used to the technology cannot oversee the costing structure of Additive Manufacturing (AM).

Especially those potential users / customers that are not used to the technology cannot oversee the costing structure of Additive Manufacturing (AM). To understand this structure, costs of the whole lifecycle of parts need to be considered. For this reason the project CoA²MPLy “Costing Analysis for Additive Manufacturing during Product Lifecycle” aims to understand and rate the cost drivers of the whole product lifecycle. Focusing on Metal Additive Manufacturing (MAM), this shall expand the fields of application for AM parts. The result will be an easily manageable framework that can be used by OEM’s, part suppliers and AM users [LJK12]. For this purpose, business processes have been modeled starting with powder production and ending with part recycling. A costing model for MAM has been developed to compare the costs of traditionally manufactured parts during different phases with the costs occurring within AM processes.

A costing model for MAM has been developed to compare the costs of traditionally manufactured parts during different phases with the costs occurring within AM processes. To have the knowledge about these processes shall help the designers to get a deeper understanding of their actions and how to influence part redesign.Nowadays many companies are not familiar with AM. Through the strong appearance in the media the topic got more important in the last years. Many companies think about AM as one possibility to increase their innovation potential. Unfortunately the knowledge about the capabilities and especially the limitations of the technology are widely unknown. They need to understand the economic influence factors of the technology to be able to select appropriate promising part candidates for a successful application of the AM technology in their company.

They need to understand the economic influence factors of the technology to be able to select appropriate promising part candidates for a successful application of the AM technology in their company. Most costing models published in literature are strongly focused on the pure production costs and do not show the overall benefits of the technology. Therefore, this work describes an approach to define lifecycle costs analysis for AM. The overall aim of this work is to help end users with the part selection and to be able to utilize AM in an economic successful manner as a long term production alternative in their companies.Methodology for assessing lifecycle costsThe methodology for estimating lifecycle costs in one part in a methodology for selecting appropriate parts for application in companies who consider using AM for their products. As the major aim of a company is to make profit, the costs of a product are still one of the most important factors for the decision making.

As the major aim of a company is to make profit, the costs of a product are still one of the most important factors for the decision making. Therefore, the costs of AM production need to be widely understood. As AM creates benefits along the production, the whole product lifecycle needs to be considered for properly assessing lifecycle costs. When talking about product lifecycle many different definitions occur. In this paper the definition of the “intrinsic product lifecycle” will be used. Compared to other definitions it considers the lifetime of a product from the first idea until the final disposal of the part, while others define the product lifecycle as the time a product remains on the market [PaBe07]. During the product lifetime a product causes costs. These costs occur at the side of the manufacturer as well as on the side of the customer.

During the product lifetime a product causes costs. These costs occur at the side of the manufacturer as well as on the side of the customer. While in the literature most people only discuss about purchasing and production costs, a deeper understanding for the costing structures need to be developed as AM has a deep impact on many aspects of the production chain. The work performed picks up the results of [LJM+12]. Based on last year’s results the lifecycle costing (LCC) aspects have been added and the model has been refined during usage.Figure 1: Different phases of the lifecycle costing modelThe developed additional LCC model is based on the German DIN 60300-3-3 standard “Dependability management - Part 3-3: Application guide - Life cycle costing”. Here, the mentioned calculation methods have been adapted to suit the comparison between traditionally manufactured parts and additively manufactured parts. The model consists of six major phases which are displayed in figure 1.

The model consists of six major phases which are displayed in figure 1. The model starts like most product development processes with a concept and definition phase. The major cost structure is set up by the costs for requirements definition as well as the design analysis. This phase is followed by the design and development phase. The considered costs include technical drawings, the component design, QM planning, a prototype and test phase as well as the production planning. This phase will be capable of evaluating different aspects of the product design phase.The costs of the production phase will by calculated by a model which is based on the work of Augsburg, Hopkins/Dickens (HD), Ruffo/Tuck/Hague (RTH) and Gibson/Rosen/Stucker (GRS). As a calculation method a “Time driven Activity Based Costing” (TD-ABC) approach has been taken. This approach allows the consideration of different influence factors on the basis the use of resources [CFG07].

This approach allows the consideration of different influence factors on the basis the use of resources [CFG07]. For the estimation of cost relevant processes, the process steps of the initial model have been simplified into four main processes:Preparation of the building job • Production of the building job • Manual removing of sample parts and support • Post processing to enhance material propertiesThese costs contain the aspects of material and machine costs considering the material- machine-combination as well as personnel costs, energy and consumption materials, post processing and quality control. The aspects of the different complexities of the building jobs can be taken into account through the use of different complexity factors. So far, this model is only applicable for the SLM production of metals (Ti6Al4,Al and 316L) but will be enhanced in the future regarding further metals and plastics.

So far, this model is only applicable for the SLM production of metals (Ti6Al4,Al and 316L) but will be enhanced in the future regarding further metals and plastics. One of the major challenges is the calculation of the building time as the influence of the building speed has shown to be one of the most important manufacturing factors [LJM+12].Equally some basic calculations for milling and casting operations have been included in the model to allow a basic comparison between several production possibilities. In consequence all phases need to be adapted as well to achieve comparable LCC results. The next phase is the installation phase. This phase contains the costs for the transport and the installation of the product.

The next phase is the installation phase. This phase contains the costs for the transport and the installation of the product. Here the influence of the possible influence of the use of monolithic structures through intelligent AM use can be shown.Following, the phase with the costs for usage and maintenance includes the costs for the maintenance necessary logistics support and the usage of the part. The usage-based costs in this model are dependent on the weight of a sample part and therefore follow one simple rule. Increase in weight equals an increase in the usage costs. Thus, this model is not necessarily fully applicable for all branches. The additional usage of AM for other branches can still be evaluated if the costs for usage are considered as zero. The logistic costs include costs for storage of spare parts and others. To make it more applicable for the aerospace industry, a methodology for assessing carbon efficiency through weight savings has been implemented.

To make it more applicable for the aerospace industry, a methodology for assessing carbon efficiency through weight savings has been implemented. Carbon efficiency costs of the logistics chain are not included.Figure 2: Displaying the product demand during part lifetime Source: [DoSC07]The last phase is the disposal phase. The disassembly costs as well as the residual value of the material or alternatively the disposal costs summarize the costs in this phase. As mentioned earlier the model can in general be used for several production processes. But adaption in all phases is needed and in some phases expert knowledge is necessary to gain exact results. The model needs to be adapted rigorously for different branches and application fields. Currently it strongly focuses on estimating lifecycle costs of AM-manufactured partsfor the aerospace industry.

Currently it strongly focuses on estimating lifecycle costs of AM-manufactured partsfor the aerospace industry. The adaption shall be facilitated in the future by the use of pre- defined application scenarios for certain branches and applications with specific scenario profiles as a guideline and as a starting point.Development of application scenario - EfficiencyPerforming a valid LCC analysis requires a deep knowledge of the different branches to which you need to adapt the existing models as many important factors vary. This starts with the different costing parameters in the companies and ends with the usage-based patterns of the product. If you compare the production targets of the automobile industry for instance in a formula student racing team and of an aerospace application you can easily understand that different factors are more important than others. Furthermore, the total production targets differ in a major way.

Furthermore, the total production targets differ in a major way. Therefore, it is necessary to develop different application scenarios in order to gather correct data for the LCC model. Once this has happened for a certain product or a certain industry, the developed LCC model can easily be transformed to other application scenarios. In the following paragraph an application scenario for the aerospace industry is described.In the aerospace scenario the civil aviation shall be considered. Here you can firstly divide the planes in five categories with different influencing factors. This differentiation starts with small airplanes with up to five tons and ends with long distance planes with up to 600 seats [HPP10]. Differentiations made in this sector regard e.g. the number of starts, landings and the total time of flight. The optimized part shall represent a part of the landing gear.

Differentiations made in this sector regard e.g. the number of starts, landings and the total time of flight. The optimized part shall represent a part of the landing gear. The landing gear systems have a significant impact on the total weight of a plane and therefore as well on the energy consumption [SCK04]. The single-aisle aircrafts like the Boeing 757 with around 250 seats are the fastest growing group of these five candidates. Therefore, this plane will be the basis of the further discussions. Between 1982 and 2004 approximately 1.050 planes of this type have been produced. This plane has two main landing gears so we estimate a use of two wheel carriers per plane which sums up to a need of 2100 parts over 22 years. The need of parts for the maintenance of the plane shall be discussed in the next chapter.Other specific factors of the use of a part for the aerospace industry arise out of the high pressure to reduce operation costs of the plane.

If you look at the kerosene price development for example, the price of a ton of kerosene is five times more expensive than ten years ago and costs nowadays around 1000 US$/[Inde13]. The significance of the fuel efficiency has raised so much in the last 10 years that the German Lufthansa has included the key performance indicator “fuel efficiency” in their operative company planning as the only non-monetary value [Luft12a]. The aspect of the possibility to reduce weight through the use of Additive Manufacturing has been proven regularly and will foster the usage of this technology in the future, as there is a strong connection between weight savings and the total fuel consumption of the vehicles. Lufthansa claims to spend 20 % of their total economic expenditures (7,4 bio. €) for fuel costs. There are plenty of models to calculate influence of weight savings in relation to fuel reduction. The differences in these models show the complexity of this topic.

€) for fuel costs. There are plenty of models to calculate influence of weight savings in relation to fuel reduction. The differences in these models show the complexity of this topic. The best-known comparison comes from Karl West who claims a cost reduction of 3000 US$ per year and plane regarding one kilo of saved weight [West11]. These costs vary strongly depending on the number of starts, landing and total flight kilometers. This work will further consider a publication of Lufthansa claiming that based on an average flight length of 2004 km 0,041 kg of kerosene can be saved for every kilogram of weight savings [Luft12a][Luft12b].In addition, the need for the reduction of carbon efficiency will become more important in the future. The aerospace industry has agreed upon a reduction plan of carbon emissions. The carbon emissions shall be reduced by 50 % from 2005 till 2050 [Kind13]. The reduction of one ton of kerosene is supposed to reduce the CO2 emissions by 3,12 tons [Frie07].

The carbon emissions shall be reduced by 50 % from 2005 till 2050 [Kind13]. The reduction of one ton of kerosene is supposed to reduce the CO2 emissions by 3,12 tons [Frie07]. In Europe every ton of carbon emission needs to be balanced by emission-certificates. The price of theses certificates is around 3 € at the moment while the aim of the European union was originally 30 €. Currently these certificates do not appear for the aerospace industry due to strong complaints of the industry. As this will not be a long-term agreement these costs will be considered in the further calculations.The combination of these two major aspects will play a key role in the evolving technology of AM and will help to foster the broader use of additively manufactured parts.

These two aspects play a major role in the later distribution of the total lifetime costs of the spare parts.Development of application scenario - Efficiency - supply chain and spare part logisticsFor economic reasons, the downtime of an aircraft should be as short as possible. So not only newly produced parts need to be considered but as well the parts, which will be needed in the later use of an airplane. AM can play a significant role in this sector , too. To further detail this aspect the following parts will help to understand the impact of AM on the supply chain and on the costs.It is estimated that an unexpected aircraft on ground costs 100.000€ a day [Göpf12]. Next to an appropriate scheduling in advance and an efficient maintenance work, it is mainly influenced by the availability of repair parts. This, however, includes highly complex and also a tremendous amount of specific components. The Lufthansa Technik AG e.g.

This, however, includes highly complex and also a tremendous amount of specific components. The Lufthansa Technik AG e.g. stores approximately 420.000 different spare parts [PfTr05] and SR Technics approximately 400.000, including parts worth a few US$ cents but also engines worth US$ 5 million. [Alte08] On the one hand, one should provide sufficient repair parts to ensure the aircraft´s airworthiness, while on the other hand one should consider the costs rising with the amount of stored repair parts. The worldwide stock of spare parts is expected to have a value of US$ 20- 30 billion and the annual costs for the storage of repair parts are expected to be 22 % of the acquisition value. These costs can be divided up into cost of capital, cost for warehousing, documentary, insurance and taxes and costs due to the excess of age of approval components, renewal, damage or theft.

These costs can be divided up into cost of capital, cost for warehousing, documentary, insurance and taxes and costs due to the excess of age of approval components, renewal, damage or theft. [PfTr05]Smaller warehousesLower capital costDesign changes are facilitatedtheLower inventoryProduction cost: AM vs. traditional manufacturingAM supports lightweight structuresDesigned partProbability of damage or theft of parts reducedApproval renewal necessarynotWarehouseDocumentary of stored part reducedIncreased speed of processesFunctional improvementsInsuranceTaxesNo exceeding of storage limitsCost-saving potential of AM in the context of spare part logisticBefore AM-machines replace entire warehouses, it is conceivable that chosen parts are subject to the production on demand. Technical limitations as well as economic the reasons, however, constrain number of parts that could be manufactured additively.

Technical limitations as well as economic the reasons, however, constrain number of parts that could be manufactured additively. In respect of spare parts logistic, research work would be improved if a greater insight into an aeronautic spare part warehouse could be gained and suitable parts examined.Higher flexibilityLogistical point of view„typical errors“ can be avoidedReduced delivery costBetter adherence to scheduleAoG-time can be reducedfigure 3:Cost-saving potential of AM in spare part logisticFurther detailed research will be carried out in the European project RepAIR. Regarding the challenges of spare parts logistic, this work reveals that AM has the potential to overcome most of these challenges. However, it has to be taken into account that current disadvantages as explained in the state of the art have to be addressed. Moreover, it is presented which cost savings could be achieved if AM machines would replace large aeronautic warehouses.

Moreover, it is presented which cost savings could be achieved if AM machines would replace large aeronautic warehouses. Based on the examinations in this work, future papers could pick up on every single potential cost saving and analyze the influence in greater detail. It would enable spare parts providers to understand the capabilities of AM better and thus research activities in this field could be intensified and improved. On top of that, a future maintenance- organization structure where the logistic department is replaced by a production department is worked out. Further research activities have to focus on how prospective business models including AM might look like.Sample for an Application Scenario ProfileFigure 4 shows the final application scenario profile for a landing gear part from the aerospace industry. It summarizes the above mentioned points.

It summarizes the above mentioned points. It takes some time to develop these description and it needs to be done carefully as many factors have an impact on the costs. Interviews with potential AM users will help to specify the exact needs and to be able to develop application scenarios.Figure 4: Application Scenario Profile Aerospace for landing gearMethodology to find verified AM part candidatesFinding appropriate parts for AM application has been complicated in the past as not enough knowledge is available and especially appropriate design rules are missing for the proper use of AM. Therefore, a methodology for finding appropriate part candidates needs to be developed as a first step for identifying appropriate parts. The methodology includes the application of the LCC model and needs to be performed by an AM expert.

The methodology includes the application of the LCC model and needs to be performed by an AM expert. The methodology is split up into three different phases (compare figure 5).It starts with an opening workshop to inform the potential end-user about capabilities and limitations of the technology. Different opportunities for successful AM application shall be presented in order to show opportunities and to break through existing thinking habits focused on traditional technologies, as in the past the one on one transformation of traditionally manufactured parts has not been shown to be optimal.The presentation shall be supported by application profiles similar to figure 4 focusing on technical possibilities, technical solutions and on the economical merit of the successful application of AM (compare [LJM+12] for benefits). This phase has the aim to bring ideas in the heads of the designers in order to make them think about possible solutions for their company.

This phase has the aim to bring ideas in the heads of the designers in order to make them think about possible solutions for their company. It can be seen as a kind of creativity fostering workshop. Future studies like [GEW13] can help to discuss the future potential of the technology. Therefore, adequate points for the entrance in the technology can be defined.For the second workshop the end user shall think about possible part candidates and define a list of specific applications and candidates. The part candidates will be discussed with the help of a value benefit analysis. Therefore, pre developed templates will be used regarding application / economic / and technical criteria. The pre-defined criteria can be weighted and completed by further factors important for a branch or a company. During discussion the characteristics will be developed. The top rated parts will be discussed again to find a starting point for application.

During discussion the characteristics will be developed. The top rated parts will be discussed again to find a starting point for application. The production costing model can help to give a first estimation of part costs.In the verification workshop further details of the parts will be clarified. includes discussions about This total technical changes and integration of a part its environment. Aspects like functional integration can be discussed in this phase. After the clarification of the part structure, an existing scenario profile can be adapted or a new scenario part profile can be developed for the usage of the part. Once completed, the results can help to identify the full potential of a solution. If the results of the analysis are negative for a the production certain candidate, values can be adapted again to future forecast of the technology.

If the results of the analysis are negative for a the production certain candidate, values can be adapted again to future forecast of the technology. This can help to define the optimal point to reconsider the use of the technology.the intofigure 5: Methodology for selecting AM part candidatesSample part for further calculationsAs a first sample to derive knowledge out of the developed models, a part which is used as an upright of a formula student racing car has been chosen. This upright was originally designed for milling (figure 6 left). One advantage is that it allows the view on two different branches. It is very similar to aerospace parts due to its high buy to fly ratio of 16.6 (94% scrap). On the right side of figure 6 the same part redesigned for AM can be seen. Supported by the Finite-Element-Analysis a “buy to fly ratio” of 1.25 has been achieved. Finally, the weight of the redesigned upright is reduced by 40%.

Supported by the Finite-Element-Analysis a “buy to fly ratio” of 1.25 has been achieved. Finally, the weight of the redesigned upright is reduced by 40%. Assuming a small batch production, a cost reduction of 50% compared to the traditionally manufactured part has been achieved. This was done without producing polluting waste. A disadvantage is the manual and thus very extensive design process based on the FEA data. [LJK+13] Three other design studies will help to derive branch dependent knowledge in the future.Figure 6: Redesign of an upright (left subtractive; right additive) Source: DMRC report 2012The findings so far show that there are several major aspects which need to be improved before (M)AM can finally play out the technological advantages and can thus be accepted as an opportunity for batch production by the broader public.

Adjusting current factors in the costing model based on expert survey [GEW12] showed cost reduction possibilities of over 50% for the future [LJK12]. As the machine costs are and will play the biggest part in this orchestra the building speed and the utilization rate will have to be improved for the future. At the moment, the character of the MAM still seems to have a wider optimization potential for batch production. To start with, software tools need to be developed. Preparation still consumes a significant amount of time. A major expertise is needed and an individual batch production is desired. In some trials human labor has consumed over 13% of the part costs.Discussion of Results and further costing developmentsThe calculations conducted have shown that the two largest contributors to the Lifecycle costs occur from the design phase and from the operation and maintenance phase as well as from the production phase.

This means an optimization of the operation costs has especially a major influence on the total lifecycle costs per part. The high costs for optimization have only been feasible for a higher number of parts produced. While the uses of the part in the racing team these costs have to be split on a production target of only four pieces. As the benefits from fuel savings are not as significant in this sector, the technology cannot shoe the full potential here. While the breakeven point between the optimized and the traditional design (both manufactured additively) lies around 250 pieces, the racing team scenario would need a production target of 425 parts in order to reach the breakeven point. In total, the additively manufactured parts have shown to be cheaper along the production chain. This effect will be enhanced if software tools will develop which can help to develop topology optimized parts with a lower amount of work.

This effect will be enhanced if software tools will develop which can help to develop topology optimized parts with a lower amount of work. One AM machine would be able to produce 673 parts per year if considering a utilization rate of 75%. It seems that one machine would be capable to deal with the production targets of the aerospace industry over lifetime. Considering common production targets like in the automobile industry high numbers like 2.000.000 parts would appear. This would mean that one machine would need 2971 years or 29 years if 100 machines would only produce this wheel carrier regarding current technological possibilities.In the considered scenarios the additive manufacturing of the wheelcarrier has shown to be cheaper as well in the production costs as well as in the lifecycle process. This shows that high complex parts like they are common in the aerospace industry have already the potential to be produce in series production with additive manufacturing.

This shows that high complex parts like they are common in the aerospace industry have already the potential to be produce in series production with additive manufacturing. Even the material costs were lower for the additive manufacturing as a 16 kg Block of aluminum is the starting point and takes around 6 hour for the build. In comparison the material needed for the additive manufactured part only displays 0,72 kg. There one can see that even higher material prices can be cheaper regarding the application.figure 7: Further findings during costing analysisIn the study “Thinking ahead the future of Additive Manufacturing – Innovation Roadmapping of Required Advancements” [GEW13] the authors have conducted a survey amongst 150 industry experts on the future development of additive manufacturing. The aim was to have a consolidated picture of the realization time of possible technical advancements in the fields of AM.

The aim was to have a consolidated picture of the realization time of possible technical advancements in the fields of AM. The study was divided into MAM and Selective Laser Melting powder bed processes as well as FDM Processes.figure 8: Excerpt of the study "Innovation Roadmapping of Required Advancements"As indicated in the paper [LJM+12] the build rate, the machine costs and the material costs as well as the utilization rate have been identified as most critical factors regarding AM production costs. Some potential has been seen for the optimization of pre and post processing of the building job as well [LJM+12]. The development of some of these factors (compare figure 8) has been investigated in the future study. As realistic measures, building rates of up to 120 cm3/h and build chamber sizes up to 2 m3 have been predicted feasible for 2020 (see figure 7). These strongly affect the costs of the build for the future [GEW13].

These strongly affect the costs of the build for the future [GEW13]. Assuming these possibilities and assuming a strong regression of material prices one can easily predict a higher accuracy for the technology in the future.figure 9: Spread of costs for producing the wheelcarrier today (left) and in 2020 (right)If the experts are right, in 2020 the total price of the sample part will drop by 90% in comparison to today’s manufacturing costs. As one can see in figure 8 the total distribution of the costs will change then. As the portion of the machine rate costs go down significantly other factors like material prices gain a lot more significance.

As the portion of the machine rate costs go down significantly other factors like material prices gain a lot more significance. This calculation estimates that the machine costs will not drop significantly in the next years as the performance and the quality control systems undergo a permanent development.Conclusion and OutlookAn appropriate part selection and the understanding of the costing structure during product lifecycle will help to foster the spread of AM already nowadays. The future development of the technology predicted by experts shows a major cost reduction potential in the next seven years. This will help to widen the application fields. The enlarged build volumes will help to make AM feasible for small batch productions and increase the efficiency. Furthermore, the larger building volumes in combination with higher build rates widen the application fields to larger parts of different industries as well.

Furthermore, the larger building volumes in combination with higher build rates widen the application fields to larger parts of different industries as well. Large parts offer even higher weight reduction potential for the lifecycle. This is important to achieve the ambitious efforts in reaching the described carbon efficiency goals. The build rate increase will make the technology by far more competitive with other technologies and broaden the application fields to areas which are not feasible nowadays. The sample part has shown that different design approaches lead to different results in lifecycle costing and production prices. The direct conversion of topology optimized into STL or in the future AMF parts will save a huge amount of money in the concept phase and make it more attractive to use. Then, the technology will be able to easily outperform traditional manufactured parts in some areas.

Then, the technology will be able to easily outperform traditional manufactured parts in some areas. Nowadays the regression effect of the optimization regarding small production numbers limits the use of the topology optimization in a major way. The knowledge of appropriate design rules (currently developed in the project “Direct Manufacturing Design Rules”) incombination with available material data and fresh educated engineers will strengthen the trust in the technology and therefore foster the further spread of AM.The presented model only represents theoretical aspects and has been a result of intense literature research to complete all data for the application scenario profiles. The methodology will be further developed as a web based tool for costing calculation and part comparison. The conversion into a proper IT-system will help to reduce the manual input in the system and allow a more detailed calculation. The whole calculation tool is constantly under refinement.

The conversion into a proper IT-system will help to reduce the manual input in the system and allow a more detailed calculation. The whole calculation tool is constantly under refinement. The further compatibility of build rates is one important topic for the future as well. Currently the build rates of different machine manufacturers are not comparable at all. Especially as the build rate is a factor which is especially dependent on the machine- material- and product- combination. These rates need to be made comparable and shall be published by the machine manufacturers for each machine material combination.In the EU-FP7 project “RepAIR”, “Future RepAIR and Maintenance for Aerospace industry Onsite maintenance and repair of aircraft by integrated direct digital manufacturing of spare parts” the part selection methodology and the assessment of real cost for MRO processes will be researched.

Furthermore, different business models will be developed and compared monetarily to find out about the total impact of AM on the supply chain and on different production models. These investigations will be performed in parallel for the opportunities of AM to redesign and to produce parts for new satellite structures. Then, the influence of functional integration will be part of the research, too.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceIMPACT ENERGY ABSORPTION ABILITY OF THERMOPLASTIC POLYURETHANE (TPU) CELLULAR STRUCTURES FABRICATED VIA POWDER BED FUSIONY. Wu*, J. Hermes†, L. Verbelen††, and L. Yang*Department of Industrial Engineering, University of Louisville, Louisville, KY 40292 †Department of Mechanical Engineering, University of Louisville, Louisville, KY 40292 ††BASF 3D Printing Solutions, Heidelberg, GermanyAbstractIn this study, experimental based investigation was carried out with various cellular structure designs realized using a developmental thermoplastic polyurethane (TPU) fabricated by powder bed fusion process, in the attempt to evaluate the effectiveness of impact energy absorption design with cellular structures when combined with favorable materials.

Various cellular designs including the re-entrant auxetic, double-arrow auxetic, octet-truss, BCC, octahedral, diamond and double bow-tie were designed and evaluated. Pendulum-rebound resilience testing and drop- weight impact testing were carried out with each designs, and the effective energy absorption capabilities of these designs were compared. The results from this study provide some initial insights into the design of TPU-based cellular structures for energy absorption applications that could benefit the establishment of more comprehensive knowledge base in this area.IntroductionOne of the technically important application areas of the cellular structures is the energy absorption applications, in which the cellular structures typically act as structural buffers and absorb energy via elastic and plastic deformation, and even fracture.

Considering that in many of the energy absorption applications the energy incidents occur as energy spikes, such as the small object impact on an aircraft structures, or the collision and drop impact on a small hand-held equipment, the dynamic energy absorption characteristics of the cellular structures should be the focus of performance designs and evaluations. Typically the dynamic energy absorption event is characterized by the high strain rate of deformation of the structures resulted from either large force or acceleration/deceleration. Due to the increased significance of factors such as inertia force, strain rate sensitivity and adiabatic heating [1], the dynamic responses of the cellular structures could be significantly different from their quasi-static and low-strain rate characteristics. There exist an extensive amount of literature that investigate different aspects of the dynamic energy absorption of cellular structures.

There exist an extensive amount of literature that investigate different aspects of the dynamic energy absorption of cellular structures. Traditionally, stochastic cellular structures, often referred to as foams, of different manufacturing technologies and base materials have been investigated for their energy absorption behaviors under various dynamic compressive strain rates [2-8]. These structures typically exhibit distinct three-stage compressive stress-strain characteristics, including the initial elastic stage, the plateau stage where the average stress levels of the structures vary relatively little as the strain accumulates, and the final densification stage, where the structures are compressive sufficiently and begin to behave like solid materials [9, 10].

Under higher strain-rate dynamic compression, the stochastic cellular structures often exhibit significantly increased2230plateau stress levels, which is attributed to various factors such as entrapped gas/fluid phase [11], the microstructural characteristics [11, 12], and intrinsic material strain rate sensitivity [11]. On the other hand, due to the general lack of control of the microscopic cellular topology, the dynamic characteristics of the stochastic cellular structures are primarily dominated by the specific manufacturing process and the relative densities.Introducing cellular topology design has been considered as a promising approach to overcome the limitation of the designability of the stochastic cellular structures. Various literatures have explored the dynamic energy absorption characteristics of various cellular topology designs [13-19].

Various literatures have explored the dynamic energy absorption characteristics of various cellular topology designs [13-19]. Compared to the stochastic cellular structures, the local topology-induced effects, such as microinertia, stress concentration and heterogeneous deformation mode becomes much more significant in the dynamic characteristics with the non-stochastic cellular designs. For examples, Lee et al. investigated the dynamic and ballistic impact responses of a 304 stainless steel pyramidal truss structure, and noted that the local microinertia kinetic effect of the strut members, as well as the intrinsic material strain rate hardening play dominant roles in the early failure stage of the structures and the total energy absorption [15]. Numerous works have suggested that auxetic designs with negative Poisson’s ratios exhibit significantly higher energy absorption capabilities compared to the regular cellular structures [13, 14]. Harris et al.

Harris et al. investigated multiple 316L stainless steel cellular-thin wall hybrid cellular structures fabricated by laser powder bed fusion additive manufacturing (PBF-AM), and suggested that the hybrid designs exhibit the most significant energy absorption design potentials in the intermedium strain rate ranges where the dynamic local buckling effect is dominant [19]. Xiao et al. studied the dynamic energy absorption of Ti6Al4V rhombic dodecahedron cellular structures fabricated by electron beam PBF-AM, and noted the feasibility of unit cell-size based impact energy absorption design [17]. In the design of the energy absorption cellular structures, the multi-functionality requirements often require highly controlled cellular topology designs that are infeasible to most traditional manufacturing processes, which necessitates the use of additive manufacturing (AM) for design realization.

Extensive amount of literature have demonstrated AM cellular structures for impact energy- absorbing purpose, and the results and discussions clearly suggested the importance of considering both the cellular geometry and the material properties [17-19].Thermoplastic polyurethane (TPU) is a class of material that is utilized broadly for the manufacturing of stochastic cellular structures for energy absorption [20], and is a potentially promising material for AM energy absorption structure realization. Cellular TPU structures fabricated via material extrusion have been explored for their various mechanical properties including energy absorption, strain rate sensitivity and quasi-static mechanical strength [21-23]. In comparison, works that based on powder bed fusion AM (PBF-AM) are limited, which might be partly due to the immaturity of the powder-based AM TPU feedstock.

In comparison, works that based on powder bed fusion AM (PBF-AM) are limited, which might be partly due to the immaturity of the powder-based AM TPU feedstock. On the other hand, due to the use of preheating and the powder feedstock, PBF-AM tends to exhibit higher geometry flexibility compared to the material extrusion based AM, which is potentially advantageous for the design of multi-objective energy absorption structures. One of such application areas is the sports helmet, which requires not only energy absorption capabilities, but also other performance criteria such as breathability, rigidity and shock buffering. Much additional works are needed to systematically understand the correlations between the cellular geometry design, the material properties and the energy absorption characteristics of TPU cellular structures fabricated via PBF- AM.

Therefore, in the current work, experimental based investigation was carried out with TPU cellular structures of various cellular designs fabricated via a laser BPBF-AM process. The results were compared to each other in order to identify potential interactions between geometry and2231materials, in the attempt to provide preliminary design selection guidelines for TPU cellular structures.Cellular Design and Sample PreparationAs in theory there exist unlimited possibilities of cellular geometry designs, it is infeasible to exhaust the design variations. Therefore, a perhaps more feasible to investigate a small subset of cellular designs that represent a variety of different design scenarios. However, there are currently limited literature about the guidance of the cellular design scenarios.

However, there are currently limited literature about the guidance of the cellular design scenarios. A few of the general design rules for the cellular structures that have been discussed include the primary structural deformation mode, which is determined via Maxwell stability criterion [24, 25], the Poisson’s ratio of the structure [26, 27], the directional isotropy [28-30], and the presence of local strut buckling [31]. For example, generally a stretch-dominated cellular structures that have Maxwell stability criterion value M≥0 tend to exhibit higher elastic modulus and initial yield strength compared to the bending-dominated cellular structures (M<0) at the same relative density levels [24]. Auxetic cellular structures with negative Poison’s ratios exhibit some unusual characteristics compared to the ones with regular positive Poisson’s ratios, such as high shear modulus, high indentation resistance, and high energy absorption [32-34].

In addition, directional anisotropy of cellular structures have been considered to establish the elastic performance bounds for the designs [29, 30]. Finally, the presence of buckling deformation failure mode within the cellular structure implies potentially catastrophic local failure [31]. Inspired by these design rules, various cellular geometries were designed and investigated in this study. 7 different types of cellular structures were designed, as listed in Table 1, which include the re-entrant auxetic (RA) and the double- arrow auxetic (DA) structures that exhibit negative Poisson’s ratios but different presence of local buckling mode, the octet-truss (OT) and octahedral (OCT) structures that exhibit stretch- dominated deformation but different directional anisotropy, the BCC, diamond (D) and double bow-tie (DB) structures that exhibit bending dominated deformation but different direction anisotropy and presence of local buckling mode.

Table 1 also lists the geometry design parameters for the cellular structures, with the corresponding variable designation shown in Fig.1. Considering the fact that relative density exerts dominant effects on the overall properties of cellular structures, for each cellular topology design, it was originally intended that 2 levels of relative densities were designed, which was realized by scaling the lengths of the struts while keeping the strut cross sectional dimension consistent (1mm). However, due to the thickness of the struts, the high density design variations for both the D and DB structures were not physically feasible and were therefore excluded from further investigation. In addition, it was later found that the high density design variation of the DA structure (DA2) resulted in un-cleanable samples, therefore the DA2 was also excluded from the experimental impact studies.Two levels of relative densities, 0.15 and 0.25, were selected arbitrarily for the study.

For each type of cellular topologies, the strut length and strut angles θ (when available) were also arbitrarily selected in that these geometry parameters might not result in the optimal impact energy absorption performance of that type of cellular topology.

Re-entrant auxetic (RA)b. Octet-truss (OT)c. BCCL1L1θL1θL1θL2d. Double arrow (DA)e. Octahedral (OCT)f. Diamond (D)g. Double bow-tie (DB)Fig.1 Unit cell designs and design parametersThe cellular unit cell designs were modeled in SolidWorks, then patterned orthogonally into a cylindrical volume using Materialise Magics. The overall geometry of the sample is shown in Fig.2a. The cylindrical sandwich panel skins have thickness of 1mm and diameter of 48mm, and the overall height is 15mm. Therefore, the overall dimensions of the cellular core are Φ48mmx13mm. The design geometry parameters were determined based on the DIN53512 standard for pendulum rebound testing. The samples were fabricated by an EOS P800 system using a developmental thermoplastic polyurethane (TPU) material at BASF 3D Printing Solutions headquarter, Heidelberg, Germany.

The samples were fabricated by an EOS P800 system using a developmental thermoplastic polyurethane (TPU) material at BASF 3D Printing Solutions headquarter, Heidelberg, Germany. The fabricated samples were cleaned using the typical2233compressed air blowing method for laser sintering powder bed fusion parts. Some of the fabricated samples are shown in Fig.2b. Multiple samples (4-5) of each design variation were fabricated, which were intended for various types of mechanical testing. The samples were each measured by digital scope and caliper for weight and dimensional information. The relative densities of each samples were consequently calculated based on the reference solid material density value of 1g/cm3 provided by the company. The dimensions and weights of each types of samples are shown in Table 2.

The dimensions and weights of each types of samples are shown in Table 2. In general, all the fabricated samples exhibit generally good dimensional accuracy and consistency, but significantly higher relative densities compared to the designed ones as shown in Table 1. This is largely attributed by the significant surface sintering that occurred on the surfaces of the cellular struts in the sandwich cores, as well as the incomplete post-fabrication cleaning, which could be noticed obviously as loose powder kept falling off from the samples during the experimental studies. Although due to the proprietary nature of the developmental material no specific process parameters were disclosed, it is expected that such excessive surface partial sintering effect is introduced by the high preheating temperature.48mmm m 5 11mma. Sandwich designb.

The quasi-static compressive testing was carried out on an Instron 5569A universal testing system. A 5kN load cell was used for the testing, and the strain was calculated from the displacement of the crosshead. Each sample was compressed to 50% of the total strain before the test terminated. Only 1 arbitrarily selected sample from each type of design was tested, which aims to establish the baseline mechanical property knowledge with2234the structures. The pendulum impact testing was carried out on a Zwick 5109 pendulum resilience rebound tester following the DIN53512 testing standard. 2 samples of each design were tested. With each sample, 10 automatically controlled pendulum impacts of 0.5J incident energy and 1.98m/s incident velocity were applied through the pendulum hammer of 15mm diameter.

With each sample, 10 automatically controlled pendulum impacts of 0.5J incident energy and 1.98m/s incident velocity were applied through the pendulum hammer of 15mm diameter. As the pendulum impact test utilizes the rebound of the pendulum for the evaluation of the elastic energy dissipation characteristic of the structures, and therefore also serves to provide a comparative basis for the dynamic energy absorption investigation of the structures that involve plastic deformation and fracture damage. The drop-weight impact testing was carried out on an Instron/Dynatup 8250 impact tester. Fig.3 shows the drop-weight impact system setup. The system was equipped with accelerometer (PCM 350B04) and dynamic force transducer (PCM 200C20), which allows for the measurement and calculation of the acceleration, velocity, location and force during the impact. 1 sample from each type of design was subjected to the drop weight impact.

1 sample from each type of design was subjected to the drop weight impact. Each sample was secured on a solid substrate by double-side tapes and subjected to up to 5 free-drop impact strikes with fixed energy level of 6.17J. For each impact strike, the energy absorption/dissipation window was defined to occur between the initiation of the impact contact and the time when the impactor speed reduced to zero, i.e. the impactor stopped moving downwards, as illustrated in Fig.3b. The total amount of energy absorption (E) during this period was calculated from the force and displacement of each step as:E(cid:32)stepn(cid:166)step1(cid:167) (cid:168) (cid:169)F F (cid:14)(cid:14) i 2i1(cid:183) (cid:184) (cid:185)x i(1)where Fi is the recorded reaction force at step i, and xi is the corresponding displacement at step i.

As the data acquisition rate of both the accelerometer and the force transducer were set at 10.24kHz, each time step corresponds to 9.765625×10-5s.Drop weight impact assemblyAccelerometerImpactor speed =0ImpactorImpact distanceForce sensorSandwichSamplea. Instrumented drop weight testing setupb. calculation of the impact distanceFig.3 Drop-weight impact testingResults and discussionsThe quasi-static compressive testing stress-strain curves for different types of cellular sandwich structures are shown in Fig.4. All the structures exhibit the typical three-stage cellular structure compression characteristics, including the initial elastic compression stage, the plateau stage, and the final densification stage. On the other hand, it was observed that the elastic stages of all the structures are rather insignificant.

On the other hand, it was observed that the elastic stages of all the structures are rather insignificant. Considering the typical low elastic modulus of the TPU materials, it was speculated that the use of crosshead displacement for the evaluation of deformation strain might not be adequately sensitive for this particular study.223532.5III2IIro a. 1.5 2V) V) ~... <l'l10.5000.10.20.30.40.50.60.5Strain- RAlRA2 - - DAl0B1 - - 01 - - OCTlOCT2 - - OTl -OT2 - - BCCl -BCC2Fig.4 Stress-strain curves of different cellular structures under quasi-static compressionIn order to overcome such issue, the stress-strain curve analysis method proposed by Christensen [35] was adopted. The schematics of the treatment is shown in Fig.5. Firstly, the stress- strain curve of a structure was fitted by a 6-order polynomial function, which was then taken to calulation the 1st and 2nd order derivatives for the stress-strain functions.

Firstly, the stress- strain curve of a structure was fitted by a 6-order polynomial function, which was then taken to calulation the 1st and 2nd order derivatives for the stress-strain functions. The yield point was identified as the first local minimum of the 2nd derivative of stress-strain functions, as marked out in red arrow in Fig.5c. Utilizing the yield point, the elastic modulus was then calculated by manually selecting the original stress-strain curve segments prior to the yield point and calculate the linear fitting functions. In addition, the total energy absorption of each structure was also calculated from the force-displacement relationships of the compressive testing up to the total strain level of 0.5. After completely unloaded, none of the samples exhibited discernable permanent deformation and visible damage.Stress-strain0.60.5y = -120.9Sx6 + 11 6.48x5 + 20.093x4 • 42.111x3 + 12.335x 2 -0.7162x +0.00861st Derivative-Eq0.4~ Q.

~ 0.3~ ~ 0.2{ 2~0.12 00.20.40.6Strain0.10.20.30.40.50.60.1Stra ina. Stress-strain curveb. 1st-order derivativec. 2nd-order derivativeFig.5 Analysis of the quasi-static stress-strain curvesTable 3 shows the elastic modulus, yield strength and energy absorption performance of each types of cellular sandwich structures. Both stretch-dominated designs, the OCT and OT structures, exhibit high mechanical properties at both relative density levels, as well as total energy absorption capabilities. The RA auxetic design also exhibit high mechanical properties and energy absorption capabilities, which appears to be in good agreement with previous studies on PA122236sandwich structures fabricated by laser sintering PBF-AM [36]. Other structures generally exhibit comparatively low mechanical properties as well as energy absorption characteristics.

Other structures generally exhibit comparatively low mechanical properties as well as energy absorption characteristics. The BCC structure in particular exhibit extremely low performance at low relative density level (BCC1-1), which was also previously observed [36]. It is also worth noting that the BCC structures were among the most easily cleanable structures, as is evident from the relative density results as shown in Table 2. Therefore, it was speculated that the loose powder that were not completely removed from the interior of some of the cellular sandwich samples might contributed to some of the energy absorption capabilities.

Due to the scheduling limitation with the samples during the rebound resilience testing, the result from RA2 design was unavailable. For the rebound resilience testing, higher rebound percentage indicates less energy dissipation during the elastic deformation of the sample. From Table 7, all the low- relative density variants of each cellular design exhibit higher energy absorption capabilities. The BCC cellular designs appear to exhibit the highest energy absorption capabilities among all the structures. There does not appear to exist a correlation between the rebound resilience and quasi-2237static energy absorption. Instead, the overall resilience performance appear to be primarily driven by relative densities of the structures. As shown in Fig.7a, it appears that there exist a linear relationship between the relative density and the rebound resilience across the entire design spectrum.

As shown in Fig.7a, it appears that there exist a linear relationship between the relative density and the rebound resilience across the entire design spectrum. Such behavior appears to suggest that for the elastic rebound energy dissipation performance, the intrinsic material property plays a more dominant role, whereas the cellular topology design might be less significant. However, such speculation should be subjected to further investigation, potentially at broader range of energy levels and impact velocities. In addition, it was also noted that there exist noticeable energy absorption capability degradation for all the cellular designs, although such trend tend to be smaller with low-relative density variations. As shown in Table 7 and Fig.7b, for all the samples of different cellular designs, the energy absorption capability degradation appear to be most significant at the first 2-3 impact strikes. Afterwards, the degradation appears to be much less obvious.

Afterwards, the degradation appears to be much less obvious. The degradation rate appears to be significantly impacted by the cellular topology design as well, although more design variations are likely needed to verify such hypothesis. Furthermore, one sample (DA1) was arbitrarily selected to undergo further impacts in order to verify the presence of performance degradation. As shown in Fig.7c, the structures continue to exhibit rebound resilience degradation throughout the entire 180-impact strike process. Such behavior was again attributed to the material characteristics of the TPU material used in this study, although no further data is currently available to further investigate such phenomenon.% different 1st-10th rebound -11.79 -29.40 -8.78 -9.58 -13.22 -3.78 -10.59 -10.59 -16.67 -11.04Avg.

Due to the use of solid substrate as well as the lack of differentiation method between the impacts on the cellular sandwich samples and the impacts on the substrate, the results were subjected to scrutiny during the consequent analysis. Fig.8 shows the accelerations and response forces of the 1st impact for each types of samples. Since no measure was available for the synchronization of drop triggering time at each drop impact, the time stamps for each impact event differ from each other, which can be clearly observed from the different occurrence time of the largest impact peaks in Fig.8, which correspond to the first impacts of each type of structures. In addition, as there is no catching mechanism in the system for secondary impacts from rebounded impactor, multiple impact peaks can be observed for each sample. For the analysis of the energy absorption characteristics, only the first and largest impact event of each structure was further investigated.

For the analysis of the energy absorption characteristics, only the first and largest impact event of each structure was further investigated. It is also noted that for the BCC structures the samples at both relative density levels appear to exhibit significantly lower impact deformation resistance, and it was clearly observed during the impact testing that the impactor compressed the BCC sandwich samples to such a degree that the substrate become a significant energy absorber. It was suspected that similar phenomenon was also present with the other designs, although from both the experimental visual observation and the emitted impact sound it was clear that substrate energy absorbing of the other structures are less significant.

From the results from quasi-static testing, all the TPU cellular sandwich structures exhibit rather low elastic modulus, which implies that during the initial elastic deformation stages of the drop-weight impact the energy absorption and deceleration processes likely to have only limited contributions to the overall energy absorption behavior. As a result, the effective energy absorbing thickness/distance of the structures is reduced, with the extreme case of the BCC structures where significant amount of energy is absorbed by the substrates.2239Impact accele rat ion10080- RAl - - 01 - - OCTl ----- OT2 - - BCCl----- RA2 - - OAl - - 0B1 ----- ocn - - on -- - -- BCC26040 "' .=! ro > ~ 20 C: 2 ~ 0 "' o:;u u <(-2058006800406080Time stepa.

However, it can be observed that for most cellular designs the impact distances are around 9-10mm, regardless of the relative density levels. The only exceptions are the low-density BCC (BCC1), which exhibits low impact distance likely due to its ineffective in deformation-induced energy absorption resulted from extremely low elastic modulus, and the RA structures, which consistently exhibit low impact distances of <8mm. It was therefore speculated that for most of the cellular sandwich structures, the thickness of the2240structures are inadequate in absorbing the incident energy. The RA structure appears to be the only exception, underlying its exceptional deceleration capability. Fig.9b shows the average energy absorption of different cellular sandwich structures over 5 impact strikes.

Fig.9b shows the average energy absorption of different cellular sandwich structures over 5 impact strikes. While it is clear that the high energy absorption of the low-relative density BCC (BCC1) structure was likely attributed to the substrate energy absorption and therefore lacks comparativity, the energy absorption of the other structures should also be subjected to extra scrutiny, mainly due to the likely presence of some substrate energy absorption. With that notion in mind, both stretch-dominated structures, the OCT and OT, exhibit higher energy absorption capabilities compared to the RA structures. Fig.9c shows the average reactive force of different cellular sandwich structures over 5 impact strikes. The significantly higher reactive forces occurred with the BCC structures again suggest that rather rigid substrate-impactor impaction took place. With the other structures, higher relative densities appear to significantly reduce the reaction force levels.

With the other structures, higher relative densities appear to significantly reduce the reaction force levels. This might again be attributed to the reduced substrate-impactor contact with high-relative density designs. The RA designs also exhibit significantly lower reactive force levels. Additionally, it was observed that while the energy absorption and impact distance of each type of cellular structure vary relatively small among the strikes, the reaction forces of the low-relative density cellular designs exhibit obvious increasing trend with increasing strike numbers, as shown in Fig.9d.10.5Average impact distance2.5Average energy absorption10 E 9.5 _§_ ~ 8.5 fl 'fi 7.5 f'c I1.1I6.540Relative density groupsos: 2.3C 0'§. 2.1] r· UJ 1.71.540Relative density groupso■ RA ■ OCT ■ OT ■ BCC ■ DA ■ DB ■ D■ RA ■ OCT ■ OT ■ BCC ■ DA ■ DB ■ Da. Average impact distanceb.

2.1] r· UJ 1.71.540Relative density groupso■ RA ■ OCT ■ OT ■ BCC ■ DA ■ DB ■ D■ RA ■ OCT ■ OT ■ BCC ■ DA ■ DB ■ Da. Average impact distanceb. Average impact energy absorption5000Average reaction force (N)3300~ 2800~ ~ 2300 C 0 ·~ 1800&1300800111 _ .• I 1140Relative density groupso4000 z iooo~ . ~1000 =--=--- ==========~~=========tooo# of imp}ct strike- RAl- - RA2 - - DAl - - D81 - - Dl - - OCTl■ RA ■ OCT ■ OT ■ BCC ■ DA ■ DB • D- - OCT2 - - OTl - - - OT2 - - BCCl - - - BCC2c. Average reaction forced. Reaction force at each strikeFig.9 Drop-weight impact characteristics of different TPU cellular sandwich designsCompared to the results from similar cellular sandwich structures made of PA12 nylon material [51], there exist some discrepancies in the performance of some of the structures. For the RA designs, the structures appear to perform consistently well regardless of the material.

For the RA designs, the structures appear to perform consistently well regardless of the material. On the other hand, the OCT structure appears to exhibit enhanced total energy absorption capabilities when TPU is used as the material, albeit also exhibiting high impact reactive force levels that might be sometime undesirable for protection purpose. Such characteristic might be associated with the fact that the softer TPU material allowed for more local elastic deformation and buckling, which reduced the softening effect that typically occurs with this type of structures when more rigid material is used. However, such speculation requires further investigation. Finally, compared to the quasi-static energy absorption characteristics, a rough correlation can be observed. The three2241cellular designs that exhibit the highest quasi-static energy absorption, RA, OCT and OT, also appear to exhibit the most desirable drop-weight impact characteristics, i.e.

The three2241cellular designs that exhibit the highest quasi-static energy absorption, RA, OCT and OT, also appear to exhibit the most desirable drop-weight impact characteristics, i.e. good compromise between energy absorptions and reactive force levels.ConclusionsIn this study, the dynamic impact energy absorption characteristics of various cellular structure designs made of TPU materials were evaluated. In general, the low elastic modulus and high ductility of the TPU materials introduced significant effect to the overall behaviors of the sandwich structures. Among the different cellular designs, the re-entrant auxetic (RA) design appears to exhibit the most prominent drop-weight energy absorption capabilities. On the other hand, the material enhancement effects from the TPU material appears to be most significant for the stretch-dominated designs, including the octet-truss (OT) and octahedral (OCT) structures.

The BCC design appears to exhibit the lowest overall drop-weight energy absorption capabilities. It is hypothesized that the sparsity of the strut arrangement and the low elastic modulus of the BCC structure might be the primary contributing factors. On the other hand, the elastic impact energy absorption of the TPU cellular designs appear to be primarily dependent on the relative densities based on the observations of the current study. Such conclusion contradicts with the general notion about the effect of geometry designs, therefore further investigations is needed.AcknowledgementThis work is partly inspired and supported by BASF 3D Printing Solutions, Heidelberg, Germany. The authors would like to acknowledge the support of experimentation from Rapid Prototyping Center (RPC) at University of Louisville.Reference [1] R. L. Sierakowski. Strain rate behavior of metals and composites. Convergno IGF XIII Cassino 1997, Italy, 1997. [2] T. Thomas, H. Mahfuz, L. A. Carlsson, K.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIMPACT OF DIRECTED ENERGY DEPOSITION PARAMETERS ON MECHANICAL DISTORTION OF LASER DEPOSITED TI-6AL-4VDavid J. Corbina, Abdalla R. Nassarc, Edward W. Reutzelc, Nathan A. Kistlerb, Allison M. Beeseb, and Pan Michalerisaa Department of Mechanical and Nuclear Engineering, The Pennsylvania State University b Department of Materials Science and Engineering, The Pennsylvania State University c Applied Research Laboratory at The Pennsylvania State UniversityAbstractThe effects of laser-based powder-fed directed energy deposition processing parameterson the distortion of deposited Ti-6Al-4V parts are assessed through in situ monitoring.

Experiments were conducted wherein substrate thickness, deposition thickness, and initial substrate temperature were varied in order to investigate their effects on distortion. Correlations of process parameters to the mechanical characteristic were also developed, uncovering some of the driving mechanisms of the measured characteristic. This work highlights the impact of substrate preheating on distortion. Most notably, the effect of initial substrate temperature on distortion depended on the size of the substrate. On thin substrates, preheating reduced the total amount of distortion. However on thick substrates, preheating increased the amount of distortion.

On thin substrates, preheating reduced the total amount of distortion. However on thick substrates, preheating increased the amount of distortion. Techniques to mitigate the unwanted mechanical defect are discussed.Keywords: Directed energy deposition; Laser cladding; Ti-6Al-4V; Additive manufacturing; Distortion; In situ measurementsIntroduction/BackgroundDirected-Energy-Deposition (DED) is an Additive Manufacturing (AM) process that canbe utilized for the repair of worn, corroded, or damaged high-value, engineered, metal components through layer-by-layer depositions involving the injection of metal powder into a melt pool formed by a laser beam.

DED provides a low-distortion repair alternative to conventional arc weld repairs for worn or damaged parts, and can be used to apply coatings with improved surface properties.The correct combination of machine process parameters is required to create a successfuldeposit, as defined by criteria related to the quality of the final build. A key criterion is the dimensional accuracy of the final part, which is affected by the temperature distribution throughout the build [1]. Unfortunately, the process of developing parameters that yield low distortion can be a time-consuming trial-and-error process. Therefore, it is essential to understand how the variation of process parameters affects distortion.1670Previous studies have investigated experimental techniques to correlate deposition process parameters to final distortion. Using strain gauges and computer numeric control (CNC) measurement techniques, Klingbeil et al.

Using strain gauges and computer numeric control (CNC) measurement techniques, Klingbeil et al. measured post-process substrate surface strains and final residual stress induced distortion that resulted from direct metal depositions of 308 stainless steel [2]. Klingbeil et al. varied substrate preheating (manually preheated using a welding torch) and deposition path, finding that preheating substrates reduces substrate warping, and that deposition paths that take advantage of inherent substrate preheating reduce warping. Jendrezejewski et al. studied the effects of base preheating on the strain-stress fields of stellite SF6 coating on X10Cr13 chromium steel, finding that preheating reduced the amount of bending and preheating at temperatures around 500°C produced crack-free coatings [3]. Deo and Michaleris studied the effect of preheating a weld region before deposition finding that preheating has the potential to achieve zero net distortion [4].

Deo and Michaleris studied the effect of preheating a weld region before deposition finding that preheating has the potential to achieve zero net distortion [4]. These studies, however insightful to the results of parameter selection on substrate distortion, do not provide insight into the real time accumulation of distortion through the process. This information can only be achieved through study of in situ measurements of distortion.Within the last decade, numerous in situ measurement techniques have been used. Plati elal. used a linear variable differential transducer (LVDT) at the free end of a cantilevered 304 stainless steel substrate to measure the bending of the substrate at a single point during the deposition of metal matrix composite (MMC) powder [5]. Ocelik et al. used 3D digital image correlation photogrammetry for in situ measurement of multiple modes of distortion of steel substrates [6].

Ocelik et al. used 3D digital image correlation photogrammetry for in situ measurement of multiple modes of distortion of steel substrates [6]. Lundbäck and Lindgren used 3D optical deformation analysis during a gas tungsten arc welding process [7].More recently studies have used laser displacement sensors (LDS) to measure substratedistortion throughout the deposition process. Denlinger et al. explored the effect interlayer dwell time has on the accumulation of distortion [8], finding that the addition of a dwell time increased the amount of distortion during the deposition of Ti-6Al-4V powder, but decreased the amount of distortion during the deposition of Inconel® 625 alloy powder. Heigel et al.

Heigel et al. investigated the effects that laser power, travel speed, and deposition patterns had on substrate distortion [9], testing longitudinal and transverse deposition patterns and found that longitudinal deposition patterns created constant accumulation of distortion with time while transverse deposition patterns produced inconsistent accumulation of distortion with time.In this study, substrate thickness, deposition thickness, and initial substrate temperature were varied between depositions of a small patch representing a possible repair procedure. The effects of the variation of substrate preheating were investigated and reported in regards to the effects they had on distortion for both thin and thick substrates.

The effects of the variation of substrate preheating were investigated and reported in regards to the effects they had on distortion for both thin and thick substrates. The goal of this work was to identify the mechanisms in directed energy deposition processes that cause distortion.2671MethodologyExperimental setupFor this study, an Optomec LENS MR-7 machine — a laser-based, DED system with a500 W IPG Yb-doped fiber laser — was used to deposit a 25.4 x 25.4 mm patch of various thicknesses of Ti-6Al-4V powder onto 76.2 x 50.8 mm (thicknesses of 2.54 or 12.7 mm) substrates of the same composition. To prevent oxidation during processing, the chamber of the machine was filled with argon, maintaining an oxygen content below 20 ppm, measured using an electrochemical sensor. During deposition, four radially-symmetrical nozzles injected Ti-6Al-4V powder at a rate of 2 g/min into the melt pool.

During deposition, four radially-symmetrical nozzles injected Ti-6Al-4V powder at a rate of 2 g/min into the melt pool. The powder used was produced by the Plasma Rotating Electrode Process (PREP®) process, then sieved to a -100/+325 sieve range (which corresponds to powder diameters between 44 µm and 149 µm). Additionally, a nozzle coaxial to the laser beam directed an argon jet to shield the focusing optics from process contaminants.Substrates were clamped into a double-cantilevered fixture, below which was a laserdisplacement sensor as, seen in Figure 1. This cantilevered clamping configuration limited out- of-plane distortion and represented a large-area repair. The fixture was arranged so that the substrates were 9.27 mm below the powder feeding nozzles, as shown in Figure 2. At the substrate surface, the second-moment laser beam diameter was 0.99 mm, as measured by a PRIMES® Focus Monitor system and the powder distribution focus was below the substrate surface.

At the substrate surface, the second-moment laser beam diameter was 0.99 mm, as measured by a PRIMES® Focus Monitor system and the powder distribution focus was below the substrate surface. During deposition, the fixture moved in the x-y direction beneath the laser via a two- axis, x-y stage, while a one-axis, z stage moved the laser processing head vertically between layers.Figure 1 Fixture used to clamp substrates3672Figure 2 Schematic of substrate beneath laser headProcess parametersTo assess their influence on deposit distortion, substrate thickness, deposition thickness,and initial substrate temperature, respectively, with all other process parameters held constant, were varied. Each of these parameters was varied between two levels, determined through preliminary testing to produce sound depositions (Table 1).

Each of these parameters was varied between two levels, determined through preliminary testing to produce sound depositions (Table 1). This paper focuses on influence of initial substrate temperature on depositions atop thin and thick substrates.Parameters Substrate Thickness Deposition Thickness Initial Substrate TemperatureTable 1 Levels of the varied process parameters Levels2.54 mm 0.76 mm ~25°C12.7 mm 2.54 mm ~400°CLaser Power [W]300Table 2 Constant process parameters Powder Flow Rate [g/min] 2Travel Speed [mm/s] 10.58Hatch Spacing [mm] 0.714Layer Height [mm] 0.254Two different substrate thicknesses were used to adjust the flexural rigidity of the substrate (Figure 3a and Figure 3b). Deposition thickness was varied by modifying the number of deposited layers. To vary initial substrate temperature, substrates were preheated using a custom-made block heater.

Deposition thickness was varied by modifying the number of deposited layers. To vary initial substrate temperature, substrates were preheated using a custom-made block heater. The heater was fabricated from a 12.7 mm thick block of steel with two evenly spaced 3/8” holes drilled through the side for heating cartridges to fit into. One 3/8”4673hole was drilled down through the center of the top surface for a thermocouple probe. An Omega® CNI8DH temperature controller was used to monitor the thermocouple probe and control the heater temperature to a set value of 400°C. The assembly was placed on top of the substrate such that the steel block was flush against the substrate and the probe touched only the surface of the substrate. The setup is schematically represented in Figure 4. During deposition, each patch was produced using the processing parameters shown in Table 2.

The setup is schematically represented in Figure 4. During deposition, each patch was produced using the processing parameters shown in Table 2. A zig-zag, parallel hatching strategy was used, as illustrated in Figure 5.Figure 3 Schematic of the combination of substrate thicknesses and deposit heightsFigure 4 Schematic of heater placement on substrate5674Figure 5 Parallel hatching used for patch build-up.Data collectionDuring deposition, the substrate distortion was measured and recorded using a KeyenceLK-031 LDS positioned directly beneath the substrate. The sensor measured and recorded the bending distortion of the substrate during deposition with a resolution of 1 μm and linear accuracy of ±1 μm.

The sensor measured and recorded the bending distortion of the substrate during deposition with a resolution of 1 μm and linear accuracy of ±1 μm. Distortion data was recorded at 20 Hz starting from the beginning of the deposition process, through the completion of deposition, and until the substrate surface cooled to 60°C to measure the full range of distortion.Omega GG-K-30 type K thermocouples laser welded to the top surface of the substrate atvarious locations were used to measure location-specific temperature history of the substrate during deposition. These thermocouples recorded temperature history at a rate of 60 Hz with a measurement uncertainty of 2.2°C or 0.75%. Temperature measurements were recorded synchronously with the timestamp, position of the laser head, and distortion measurements.

Temperature measurements were recorded synchronously with the timestamp, position of the laser head, and distortion measurements. The locations of the welded thermocouples on the surface of the substrate along with the location of the LDS measurement are illustrated in the schematic in Figure 6.6675Figure 6 Schematic of the top (a) and bottom (b) of the substrate illustrating the location of the measurement sensorsAfter deposition, the patch height was measured on from each run to ensure proper buildsize. The height was measured at the center of the patch using digital calipers.Results and DiscussionAdjustments of processing parameters resulted in variations in in-situ measurements ofthermal deformation.

The effects of substrate preheating on distortion were analyzed by isolating the change in preheat between the results of four separate runs and comparing the corresponding in situ distortion measurements.The results showed that preheating of a substrate before deposition had an impact on the amount of total distortion and the rate of distortion. When preheating a thin substrate (2.54 mm thick), the amount of distortion decreased by 22.1%, from 0.83 mm to 0.65 mm, relative to the room-temperature case (Figure 7 (left)). By the end of the first layer, the effect of preheating had already decreased distortion by 0.20 mm (33.3%), and by the end of the second layer it had decreased distortion by 0.20 mm (30.4%). By the end of the third layer distortion had decreased by 0.20 mm (24.3%).When preheating a thick substrate (12.7 mm thick), the amount of distortion was shownto increase (Figure 7 (right) and Figure 8).

The final measured increase in distortion was 111.9%, relative to the room-temperature case. Distortion of the preheated case did not reach a steady- state before the end of data collection. The distortion results of the preheated 12.7 mm thick substrate also exhibited similar distortion rate during the deposition of each of the 10 layers. Most distortion occurred after deposition concluded, during cool-down.7676Figure 7 Effect of substrate preheating on distortion for a 0.76 mm thick deposit on a 2.54 mm thick substrate (left) and on a 2.54 mm thick deposit on a 12.7 mm thick substrate (right)Figure 8 Effect of substrate preheating on distortion for a 2.54 mm thick deposit on a 12.7 mm thick substrateThe most notable result was that substrate preheating had differing effects on distortion depending on the thickness of the substrate. Preheating of a thin substrate decreased the amount of distortion whereas preheating a thicker substrate increased the amount of distortion.

Preheating of a thin substrate decreased the amount of distortion whereas preheating a thicker substrate increased the amount of distortion. By preheating a thin substrate, the temperature difference between the substrate surface and the melt pool was reduced. In the case of the preheated substrate, the higher surface temperature will lead to a reduction in the thermal gradient, which should result in a reduction of distortion. But this effect is not observed with the thick substrates.8677These counter-intuitive results could be attributed to a variety of reasons. One possibleexplanation could relate to the method of preheating. In this study, substrate preheating induces a localized heat source to only the top surface of the repair substrate. With this method, a thin substrate would be heated to a more uniform temperature before deposition, whereas a thick substrate would have a larger temperature difference between the top and bottom surfaces of the substrate.

These temperature differences could indicate a difference in thermal gradient. The actual reasons behind the counter-intuitive effect of thick-substrate preheating are under investigation.ConclusionsThis study presented an analysis of in situ distortion data from a series of laser-deposited Ti-6Al-4V patches wherein substrate thickness and preheating were varied. It was shown that the distortion of the thin geometries can be mitigated by increasing the initial substrate temperature. However, preheating of thick substrates increased substrate distortion. The reasons for this effect are unclear and warrant further investigation.

However, preheating of thick substrates increased substrate distortion. The reasons for this effect are unclear and warrant further investigation. Although no quantitative analysis of the thickness of the substrate was performed, the results of this experiment suggest to avoid preheating of thicker substrates or to perform preheating on a thinner substrate to mitigate any effects of part distortion during the deposition process.AcknowledgementsFunding: "This material is based on research sponsored by the Air Force ResearchLaboratory through America Makes under agreement number FA8650-12-2-7230. The U.S. Government is authorized to reproduce and distribute reprints for Governmental purposes notwithstanding any copyright notation thereon."References[1] Mukherjee T, Zuback JS, De A, DebRoy T. Printability of alloys for additivemanufacturing. Sci Rep 2016;6:19717. doi:10.1038/srep19717.[2] Klingbeil NW, Beuth JL, Chin RK, Amon CH.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperImpact of Embedding Cavity Design on Thermal History between Layers in Polymer Material Extrusion Additive ManufacturingSwapnil Sinha1 and Nicholas A. Meisel21Department of Mechanical Engineering, Penn State, University Park, PA 16802 2School of Engineering Design, Technology, and Professional Programs, Penn State, University Park, PA 16802AbstractBy pausing an additive manufacturing process in mid-print, it is possible to create multifunctional structures through strategic insertion of foreign components. However, in polymer material extrusion, previous research has shown that pausing the build decreases the eventual strength of the final part, due to cooling between layers.

However, in polymer material extrusion, previous research has shown that pausing the build decreases the eventual strength of the final part, due to cooling between layers. To better predict this part weakness, this paper seeks to quantify how the toolpath affects the thermal history of a cross-section, thus impacting the formation of weld strength between printed layers. This is pertinent to in-situ embedding as different embedded geometries will require different cavity designs, which, in turn, will affect toolpath design. In-situ thermal measurements are experimentally collected with a thermocouple at the layer interface of structures with different cavity designs. The weld strength between layers is then obtained through tensile tests and theoretically evaluated using polymer weld theory. Results show more accurate predictions of load at failure with this method.1.

Results show more accurate predictions of load at failure with this method.1. IntroductionThe layer-by-layer nature of Additive Manufacturing (AM) provides an opportunity to embed sensors, actuators, circuits, and other functional components within a part, during the manufacturing process [1,2]. For in-situ embedding with the material extrusion process, first, a cavity is designed in the part and the printing process is paused once the cavity is printed. The foreign component is inserted into the cavity, and the print is resumed. This integrates the functionality of the inserted components into structures along with the offered flexibility in viable designs. This capability recently has been recognized for its application to the design and production of optimized electronics in aerospace industries and sophisticated “smart” objects [3– 7].

This capability recently has been recognized for its application to the design and production of optimized electronics in aerospace industries and sophisticated “smart” objects [3– 7]. The flexibility in manufacturing offered by AM allows cost effective manufacturing of customizable embedded electronics for applications where weight and volume is critical to overall performance. Various assembly-related processes and resource waste are eliminated with in-situ embedding, which makes AM a vital tool contributing towards the vision of Industry 4.0 to revolutionize manufacturing.While in-situ embedding with AM shows potential for the future of manufacturing, its layer- by-layer process also presents with some limitations. AM structures are specifically weaker in the build direction [2,8]. Furthermore, cavities must be designed with orientations that allow foreign components to be completely submerged into the part to avoid any damage to the print head [9].

Furthermore, cavities must be designed with orientations that allow foreign components to be completely submerged into the part to avoid any damage to the print head [9]. However, prior research has shown that the required process interruption for embedding process directly reduces the tensile strength of printed parts [10]. The introduced weakness at the paused layer, as well as further reduction of part strength due to the designed embedding cavity, raises concern over direct manufacturing of products via in-situ embedding with AM.1778To address these weaknesses due to embedding, understanding how various factors in the process directly influence the part strength is important. One such factor in the material extrusion process is the thermal history at the interface of deposited material, that has been known to directly control the part strength [11,12].

One such factor in the material extrusion process is the thermal history at the interface of deposited material, that has been known to directly control the part strength [11,12]. This thermal history is subject to change with the material, ambient temperature, printing parameters like speed or layer height, and the material deposition toolpath [12–14]. Previous studies have almost exclusively focused on either the impact of these process parameters on the strength of the AM structures or the material’s behavior under the processing temperature. To build this gap in design and theory of materials, this research focuses on investigating effects of different cavity designs for embedding on thermal history between the layers, especially due to different toolpath of material deposition. For this purpose, printed structures with different embedding cavity designs are studied with a computerized tomography (CT) scans, and the cross section area is categorized on the basis of the toolpath.

Thermal measurements are obtained for each of these categories and polymer weld theory is utilized for prediction of strength. The predictions are then validated through uniaxial tensile tests.2. Literature ReviewThe following sub-sections summarize literature of relevance to the in-situ embedding process with material extrusion AM. The next section details the design for embedding considerations. Section 2.2 describes the link between toolpath and temperature of the layer interface. Section 2.3 details the polymer theory for predicting strength.2.1 Design for in-situ embeddingAs with other manufacturing processes, a successful AM part with embedded components requires certain design considerations before manufacturing. To avoid damage to the print head or the embedded component, cavities must be designed so that the component is completely submerged and secured into the print [15–17].

To avoid damage to the print head or the embedded component, cavities must be designed so that the component is completely submerged and secured into the print [15–17]. Direct writing techniques to manufacture electronics also require process-based design planning such that the wire interconnects are accurately formed [18–21]. While these literatures focus on exploring applications and development of hybrid processes for embedding, few studies address the design for embedding considerations needed to retain the structural strength of the printed part [22].Research suggests that careful design of cavities is essential for viable in-situ embedding. However, these cavities can also act as defects and directly affect the part’s strength. Furthermore, the presence of cavities also influences the toolpath for material deposition. For example, in a material extrusion process, each layer starts by printing perimeters or shells.

Furthermore, the presence of cavities also influences the toolpath for material deposition. For example, in a material extrusion process, each layer starts by printing perimeters or shells. These shells define the boundary of the layer, according to the shape of that layer (See Figure 1). Research has shown that the presence of these shells is essential for the final quality of the part and directly influences the finish and strength of the structures [23–25]. The layers with the embedding cavities have additional perimeters defining the boundary of the cavity itself. Next, the empty spaces inside the perimeters are filled with the infill, for which users can typically select both density and pattern. A large body of literature exists that centers on optimizing the toolpath to minimize the time to print a layer [26–28]. Few address toolpath optimization based on structural strength [29].

A large body of literature exists that centers on optimizing the toolpath to minimize the time to print a layer [26–28]. Few address toolpath optimization based on structural strength [29]. However, for embedding, the strength between layers is of particular interest as the process is typically must be paused between the layers. This process interruption has been shown to directly1779weaken the structure at the paused layer in prior research [10,22]. This strength is dictated by the thermal history of the interface.(a)(b)FIGURE 1. Shows the digital layer display in the Cura software, for a print layer. (a) Shows the toolpath for a 100% dense print with rectangular pattern infill. (b) Shows a toolpath for 20% dense infill for a triangular pattern infill.2.2 Toolpath and thermal historyIn order to increase the interfacial strength between layers for in-situ embedding AM, the thermal history, and thus the toolpath, must be considered.

A typical desktop polymer material extrusion process uses polymer filament material that is fed in the nozzle, melted, and deposited layer by layer [30]. As the molten polymer is deposited in its viscous state, it cools and solidifies into roads of elliptical cross section due to the shear force on material by the nozzle [12–14]. This also leads to inbuilt air gaps between the road and the layers, which impact the overall strength of the structure [31–33]. The final strength of the part depends on the thermal history at the interface of the deposited and depositing roads; this thermal history defines the weld strength at the layer interfaces [12,14,34,35].The thermal history at the layer interface depends on the toolpath, since the toolpath defines the location of molten material over time [11,14]. As discussed in Section 2.1, the toolpath varies in different regions and thus varies the thermal history at the layer interface.

As discussed in Section 2.1, the toolpath varies in different regions and thus varies the thermal history at the layer interface. This results in differences in weld strength at each point of the embedding cross section [36,37]. A series of literature has indicated that all the print parameters that directly influences part strength can be attributed to the thermal history at the material interface [11,12]. To predict how this thermal history leads to weld formation in printed parts, it is necessary to understand the polymer’s unique response to temperature quantified through its rheology [35].2.3 Prediction of weld formation in polymer materialThermal bonding or weld formation at the interface depends on material in processing and how cohesion occurs in that material. A polymer’s behavior with temperature is unique because the long polymer chain molecules exhibit random motion at higher temperatures.

A polymer’s behavior with temperature is unique because the long polymer chain molecules exhibit random motion at higher temperatures. This motion is known as reptation and causes polymer molecules to entangle with each other [38]. These entanglements result in cohesion at the interfaces. However, polymer molecule reptation is highly dependent on the time for which it stays at any elevated temperature. In isothermal states, the degree of cohesion can be directly measured with Ezekoye equation [34], using knowledge of the polymer’s reptation time. Reptation time is the time taken by the polymer chain to completely1780move out of its initial boundary at a given temperature [34].

The interface between the polymer material will heal to 100% and exhibit the bulk-strength if it stays at a high temperature for the duration of reptation time for that temperature. Wool and O’Connor [38] used this equation to estimate the degree of healing in a polymer interface in isothermal conditions. Bartolai et al. [8] used equation 1 to develop a strength predicting “weld theory” for the non-isothermal case of material extrusion AM, which requires reptation time as a function of temperature. They demonstrated that by collecting thermal history at the road interfaces, and incorporating air gaps in the cross-sectional area measurements, the model was successful in predicting strength.

For each subsequent time step in the thermal history, the obtained (cid:2026)(cid:3050)(cid:3032)(cid:3039)(cid:3031)(cid:3281) is used as the new (cid:2026)(cid:3050)(cid:3032)(cid:3039)(cid:3031)(cid:3284). Bartolai et al. demonstrated the applicability of equation 2 in predicting the weld strength along the X-Y direction of the build. However, for in-situ embedding, the process must be paused between the layers. Therefore, it is necessary to adapt the previous weld theory to predict the strength along the Z direction of the build.3. Materials and MethodsThe aim of this paper is to use polymer weld theory to predict weld strength between layers for printed specimens with different toolpath designs. This will enable an understanding of how the cavity design affects the strength of the structures. As discussed, the strength of welding between the layers is directly correlated with the thermal history at the interface [11,37].

As discussed, the strength of welding between the layers is directly correlated with the thermal history at the interface [11,37]. Therefore, thermal history measurements of the layer interfaces were first obtained. CT scans of the printed specimens were then taken to obtain the actual contact area of the embedding layer’s cross section. The cross sections were divided into zones of similar thermal history based on the type of toolpath (i.e., boundary, infill). To use the polymer weld theory for strength prediction, material rheology measurements were also made. Finally, to validate the strength predictions, experimental specimens were tensile tested until failure. The following sections discuss these methods in detail.3.1 Specimen design and preparationTo analyze the strength between layers, a series of tensile specimens were designed and printed upright, with their length along the Z direction of the build plate.

However, ASTM standard tensile1781(1)(2)specimens require smaller cross section areas than necessary to accommodate cavities for embedding. Therefore, a rectangular specimen of cross section 24 mm (cid:3400) 12 mm and length 96 mm was specially designed for the purpose of tensile tests. A series of control specimens were prepared without an embedding cavity; these then serve as the baseline for comparison. In order to make the results of this research broadly applicable, basic primitive geometries (e.g., cuboid, triangular prism, and cylinder) were chosen for the initial cavity designs. Each printed tensile specimen was designed with the cavity located at the geometric center of the specimen. Groups of samples with four different cuboid orientations (denoted as Cu1, Cu2, Cu3, and Cu4, as seen in Figures 2a, 2b, 2c, and 2d, respectively) were prepared.

Groups of samples with four different cuboid orientations (denoted as Cu1, Cu2, Cu3, and Cu4, as seen in Figures 2a, 2b, 2c, and 2d, respectively) were prepared. It is crucial to note that, each different orientation results in a different cavity cross-sectional area at the specimen’s embedding plane. Additionally, a group of samples with an equilateral triangular prism cavity (denoted as Prism, Figure 2e) and a group with a cylindrical cavity (denoted as Cyl, Figure 2f) were prepared.All build preparation and manufacturing parameters were kept consistent with a layer height of 0.2 mm and 100% infill density, prepared with Cura-lulzbot-3.2.27 software. The toolpath was automatically generated in the Cura software; this results in each specimen groups exhibiting a different toolpath due to differences in cavity cross-section.

The toolpath was automatically generated in the Cura software; this results in each specimen groups exhibiting a different toolpath due to differences in cavity cross-section. The natural PRO Series PLA (Polylactic acid) 3 mm filament by MatterHackers™ was used to print the specimen with a desktop material extrusion system (Lulzbot Taz 6). Specimens were printed one at a time, oriented with their length (96 mm) along the Z axis and their base centered on the build plate. An MTS 100 system with a 3000 lbf load cell was used to conduct uniaxial tensile tests until the specimens failed.

Unfortunately, capturing this comprehensive thermal history experimentally is challenging due to obstruction from the motion of the printing nozzle. Therefore, to simplify measurement, the cross section of each specimen was divided into three zones: shell, infill, and infill boundary (see Figure 3). The shell zone occurs at the external perimeter of the cross section (shown in orange in Figure 3), while the infill zone is the central area of the cross-section filled with alternating roads of material (shown in green in Figure 3). Finally, the infill boundary zone (shown in blue in Figure 3) occurs where the infill meets the shells. In this zone, the tool changes direction, which causes it to decelerate, stop, and then accelerate. This results in a concentration of thermal energy as well as thicker material deposition [8].■ Shell ■ Inflll ■ Infill BoundaryBoundary FIGURE 3. The shell, infill, and the infill boundary zones in the cross section of a control specimen.

The shell, infill, and the infill boundary zones in the cross section of a control specimen. The air gaps are visible as white spaces between the deposited roads in the CT scanned slice of the cross section (left)As shown in Figure 3, the deposition of polymer roads naturally results in air gaps present in each layer. As such, to obtain the actual cross-sectional area necessary for weld strength calculations, samples were CT scanned with a GE v|tome|x L300 multi-scale nano/micro CT system at a 16-micron resolution. The software program Image J was used for analysis of the obtained scans. The image slices were converted to binary, representing scanned point cloud data as black (0) or white (255) for analysis. The CT scanned slices were sectioned into shell, infill, and infill boundary zones and the area of each zone was calculated. Three slices were assessed for each group by counting the number of colored pixels in each zone using Image J.

Three slices were assessed for each group by counting the number of colored pixels in each zone using Image J. For the infill zone, this calculation includes overhanging roads that are not in contact between the layers due to the changing infill orientation in each layer. To compensate for overestimation of infill contact area, the computed cross-section area of each infill zone was reduced by half for future calculations.To experimentally determine the thermal history of each of the three toolpath zones, a thermocouple was used to capture the temperature-time data for each zone during printing. A type J thermocouple (KK-J-245) by OMEGA Engineering was used to obtain the thermal history at the interface (see Figure 4b and 4c) through a DI-245 by DATAQ instruments data acquisition system.

A type J thermocouple (KK-J-245) by OMEGA Engineering was used to obtain the thermal history at the interface (see Figure 4b and 4c) through a DI-245 by DATAQ instruments data acquisition system. Holes were designed in the control specimen to allow the thermocouple wires to be embedded.1783The print was paused once the holes were printed, then the two thermocouple wires were inserted and brought in contact at the paused layer. Once the build is resumed, the thermocouple should capture the thermal history at the interface between the paused and the resumed layer (Figure 4d). The thermal histories for the three zones were obtained by carefully positioning the thermocouple contact point in these zones for a control specimen. Two measurements for each of the zones were repeated for consistency in measurements.Fresh layer (T.)Interface ____._[T;=Avg(T.,Td)] /Deposited layer (Td)Build plate(a)(b)(c)(d)FIGURE 4.

(a) Snapshot from the IR camera while recording the thermal history. (b) Embedded thermocouple wires in red and green, through the designed cavity, connected at the exposed paused layer. (c) Shows the location of thermocouple at the interface of deposited and resumed layer. (d) Shows the representation of the print, and the layer interface location with respect to the deposited and the fresh layer.When pausing the print to place the thermocouple wires in the specimen, the topmost layer will naturally cool down more than would be observed in continuous print. This may skew calculation of the average thermal history between layers. Because of this, an IR camera (Optris Pi 450, by Optris Gmbh, with Optris Pi Connect thermal imaging software) was also used to collect the temperature-time data at the topmost visible layer of a continuous printing process (see Figure 4a).

The temperature data from the IR camera shows that, in a continuous print, the previously deposited layer has already cooled to ~60 °C at the moment that a fresh layer is deposited. However, the necessary pause to embed the thermocouple causes the previously deposited layer to cool down to ~35 °C at the moment a new layer is deposited (a discrepancy of ~25 °C). Assuming that the temperature at the interface is an average of the temperature of the paused layer and the freshly deposited layer (Figure 4d), the interface temperature history for a continuous print can be estimated by adding 12.5 °C to the interface temperature obtained from the thermocouple.3.3 Polymer material rheology for weld strengthof the material as a function of temperature is required. However, as discussed in Section 2.3, the strength prediction from polymer weld theory is only valid for temperatures at which the polymer molecules reptate. Previously, Bartolai et. al.

However, as discussed in Section 2.3, the strength prediction from polymer weld theory is only valid for temperatures at which the polymer molecules reptate. Previously, Bartolai et. al. used the temperature history only above the Tg value for both Polycarbonate (PC) and Acrylonitrile butadiene styrene (ABS); this temperature allows the weld forming polymer molecule movements to occur [8,14]. While both PC and ABS materials are amorphous in nature, exhibiting a clear Tg region, the PLA material used in this research is semi- crystalline. The partial crystalline arrangement of its molecules inhibits their movement until its melting temperature is reached [39], which suggests that PLA molecules will only diffuse and weld when the polymer is at its molten state [40,41].

Therefore, to predict the weld strengthTo calculate the weld strength from Equation 2, (cid:2028)(cid:1191)(cid:3045)(cid:3032)(cid:3043)1784formation at the layer interface, only the thermal history for the polymer is at its molten state is required. To obtain the melting temperature for PLA, Digital Scanning Calorimetry (DSC) experiments were performed for a heat-cool-heat cycle, where the material was heated from 25 °C to 170 °C at a rate of 10°C per minute. The melting peak occurred at approximately 148.6 °C for both heat cycles, with the end of the peak occurring at 150 °C. Since the reptation of molecules that causes adhesion occurs only in PLA’s molten state, only the temperature history of the fresh layer above 150°C is of interest for predicting the strength. Therefore, only the temperature values above 150 °C in the thermal history of the fresh layer was considered in this research.

Therefore, only the temperature values above 150 °C in the thermal history of the fresh layer was considered in this research. Accordingly, the interface temperature values (which is assumed to be an average of the temperature values of the fresh layer and the deposited layer), resulting in predicted fresh layer temperature above or equal to 150 °C were considered for prediction.After identifying the temperature region of interest for PLA, material rheology measurements were made. Hot pressed PLA disks were prepared and storage and loss modulus were obtained using a strain-controlled ARES rotational rheometer. Measurements were performed at angular frequency range of 0.5 rad/s to 100 rad/s, at strain rate of 1% for every 10 °C from 160 °C to 200 °C.

Measurements were performed at angular frequency range of 0.5 rad/s to 100 rad/s, at strain rate of 1% for every 10 °C from 160 °C to 200 °C. For viscoelastic materials, these isothermal sweeps do not change in shape but only shift along the time axis with temperature (see Figure 5a), which makes it possible to obtain a master curve at a desired reference temperature (Tref, see Figure 5b). This phenomenon, known as the time- temperature superposition (TTS) principle, uses a temperature-dependent translation factor, aT to perform a shift operation on the master curve. This enabled calculation of the isothermal modulus properties for any other temperature.

., 0 Q) E ., e- 180_0 "' + 190_0 m 0 +200_0~_J10 510•10310210 110 110•10 110210•10 1Angular frequency w (rad/s)Angular frequency w (rad/s)FIGURE 5. (a) Isothermal time-temperature superposition (TTS) of storage modulus and loss modulus for PLA, and (b) TTS master curve at reference temperature 180 °C.4. Results and AnalysisThis section uses the methods from Section 3 to predict the tensile load at failure for specimens with different in-situ embedding cavities. This prediction leverages both contact area and thermal history at layer interfaces. To validate the accuracy of the prediction, experimental tensile tests are also conducted.4.1 Impact of cavity design on thermal history at the interfaceThe cross-sectional area was measured through the CT scanned images of the interfaces, as detailed in Section 3.2.

The total cross-sectional area for each specimen lies in the range of 164 to 185 mm2, but the cross-sectional area distribution between the zone types change notably depending on cavity geometry. CT data, summarized in Table 1, shows that, since the presence of a cavity introduces more perimeters, more cross-sectional area becomes dedicated to the shell zone (see Figure 6.).■ SbeUCT scanned slice of the cross section■ Infill■ InfillBoundaryControlCu_ JCu_2Cu_JC u_4PrismCylFIGURE 6. CT scanned slice and digital slice of the specimen.1786102TABLE 1.

CT scanned slice and digital slice of the specimen.1786102TABLE 1. Cross section area distribution within the different samples.GroupsControl Cu1 Cu2 Cu3 Cu4 Prism CylArea (mm2) 70.9 83.1 83.5 75.8 79.9 102.3 93.4Shell% of total 39% 51% 46% 41% 48% 55% 51%Area (mm2) 61.3 32.7 40.8 36.0 35.6 31.3 30.0Infill% of total 33% 20% 23% 19% 21% 17% 16%Infill Boundary Area (mm2) 50.8 48.2 55.8 73.7 51.0 51.3 58.3% of total 28% 29% 31% 40% 31% 28% 32%Total c/s area (mm2)183.0 164.0 180.0 185.5 166.5 184.9 181.7For each of the three zones, the thermal history was obtained by carefully positioning the thermocouple wires in the control specimen’s cross section, as discussed in Section 3.2. The resultant temperature history for each zone is shown in Figure 7. In the plot, abrupt jumps in temperature denote the deposition of molten material in close proximity to the thermocouple.

2 "' ... Q) a. 8 60 ~040201 layer00246810 Time (s)1214161820FIGURE 7. Thermal history as obtained by the thermocouple for each zone for one layer of material deposition with the location of measurement in the control specimen cross section.4.2 Validation failure load predictions using weld theory and rule of mixturesAs discussed, the different thermal history in each of the three cross-sectional zones leads to different weld strength formation for each of them. Therefore, to simplify prediction, these AM structures can be treated as composites made from sections with different strength values.

Similarly, As, Ai, and Aib are the actual cross-sectional areas in each of the three zones. The ultimate tensile strength ((cid:2026)(cid:3022)(cid:3021)(cid:3020)) for PLA was taken as 50 MPA [43].With the known function of (cid:2028)(cid:3045)(cid:3032)(cid:3043)(cid:4666)(cid:1846)(cid:4667) from Equation 3, the (cid:2028)(cid:1191)(cid:3045)(cid:3032)(cid:3043) for each time step from the recorded thermal history was obtained and substituted in Equation 2. The time step (tf-ti) was taken as 0.04 seconds, which is equal to the resolution of the data acquisition system for the thermocouple. Subsequent calculations for each time step where the freshly deposited layer was above the melting temperature (150 °C) gave the predicted final weld strength. Through Equations 2 and 3, the weld strength of the shell, infill, and the infill boundary zones were obtained as 27 MPa, 45.3 MPa, and 39.2 MPa, respectively.

Through Equations 2 and 3, the weld strength of the shell, infill, and the infill boundary zones were obtained as 27 MPa, 45.3 MPa, and 39.2 MPa, respectively. The failure load for specimens with different embedding cavity geometries were then predicted using Equation 4.The results of tensile tests of different embedding cavity specimens, along with their predicted strength values derived from polymer weld theory are shown in Figure 8. For comparison, the figure also shows the predicted load at failure if the cross section were not divided into zones, and only the total area was used by assuming one thermal history throughout the cross section.□ Experimental■ Predicted wit weld zones■ Predicted wit shell■ Predicted w/t infill■ Predicted w/t infill boundary1000090008000g 7000 .a ;aQ)60005000~ "Cl ce 0 i-l40003000~ MMt N MMM[M , M r* M MMMr *M.. •MM200010000ControlCu3Cu2Cu4CulPrismCylFIGURE 8.

•MM200010000ControlCu3Cu2Cu4CulPrismCylFIGURE 8. Experimentally obtained and predicted tensile load at failure for all the groups of specimens.1788(4)TABLE.2. Summary of the experimental and predicted loads at failure for the groups with errors.GroupsPredicted load distribution in the cross section (N)Tensile load at failureControl Cu3 Cu2 Cu4 Cu1 Prism CylShellInfill1915.0 2776.9 2047.6 1632.2 2253.8 1848.4 2158.4 1613.6 2243.3 1480.5 2761.9 1419.0 2521.6 1358.4Infill Boundary 1991.6 2889.1 2187.8 1999.5 1890.8 2012.5 2285.8Predicted (N) 6683.5 6568.9 6290.0 5771.5 5614.7 6193.4 6165.8Experimental Avg (N) 6991.6 5530.9 6384.3 5614.1 5400.5 5831.8 5404.0% error 5% 16% 1% 5% 4% 7% 12%Cyl and Cu3 groups show high error in strength prediction by considering the three weld zones. The experimental load at failure were lower than predicted for these two groups.

The experimental load at failure were lower than predicted for these two groups. Qualitative observations show that under extrusions in the infill zone were the reason for low strength value. Furthermore, 80% of the tensile failures for the cavity specimens occurred at the layer where the cavity started or ended printing. Stress concentrations due to the abrupt start of the cavity could explain this location of failure. The other 20% of the time, the failure occurred at layers with the cavity. With the added process interruption required for embedding, this location of failure is expected to be the paused layer due [10,22].As is evident from the Figure 6 and Table 2, the proposed method of dividing the cross section into weld zones and leveraging them to predict the load at failure through the rule of mixtures more accurately predicts the load at failure than assuming a constant thermal history throughout a printed layer.

This method results in prediction error as low as 1%, depending on the geometry of the embedding cavity. By comparison, approximating the cross section as a bulk with only one thermal history either under- or over-predicts the load at failure. Establishing this understanding is a crucial step toward enabling engineers to account for cavity geometry and toolpath design in their estimation of part strength for multi-functional embedded AM parts.5. Conclusion and Future WorkThe research presented in this paper proposes a method to predict the strength of parts with different cavity design for embedding applications with material extrusion AM. The thermal history of printed layers is used in conjunction with PLA’s rheology behavior to predict the strength between the layer interfaces using the weld theory by Bartolai et. al [8]. Interfaces were considered as composites made of zones with different thermal histories.

al [8]. Interfaces were considered as composites made of zones with different thermal histories. The experimental results confirm that this method of considering thermal history in a layer results in more accurate predictions of failure load when compared with assuming a constant thermal history across the layer. This demonstrates the importance of toolpath design when incorporating cavities to support in-situ embedding in AM.This research demonstrates that the thermal history at the layer interface as a result of the toolpath and material distribution has a considerable impact on printed part strength. However, the methods presented approximated the thermal history for each cross-sectional zone as one, due to limitations in the thermal data collection techniques. Additionally, variations in the speed of material deposition due to changes in toolpath direction were ignored.

Additionally, variations in the speed of material deposition due to changes in toolpath direction were ignored. Because of the high sensitivity of the thermal histories to the toolpath, these effects should be considered to improve1789prediction accuracy. This requires development of a computational tool to more precisely consider toolpath design and capture thermal history at each point in a layer. The thermal histories for each point can then be used to improve the accuracy of weld strength prediction between the layers.6. References [1] Kumar, V., Rajagopalan, S., Cutkosky, M. R., and Dutta, D., 1998, “Representation and Processing of Heterogeneous Objects for Solid Freeform Fabrication,” Geom. Model. Work., pp. 1–21. [2] Meisel, N. A., Elliott, A. M., and Williams, C. B., 2015, “A Procedure for Creating Actuated Joints via Embedding Shape Memory Alloys in PolyJet 3D Printing,” J. Intell. Mater. Syst. Struct., 26(12), pp. 1498–1512.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIMPACT OF FLOW AID ON THE FLOWABILITY AND COALESCENCE OFPOLYMER LASER SINTERING POWDERR.G. Kleijnen*, M. Schmid*, and K. Wegener†inspire AG, Innovation Center for Additive Manufacturing Switzerland, Lerchenfeldstrasse 3, 9014 St. Gallen, Switzerland†Swiss Federal Institute of Technology, ETH Zürich, Institute of Machine Tools andManufacturing (IWF), Leonhardstrasse 21, 8092 Zurich, SwitzerlandAbstractSmall amounts of nanometer-sized flow aids are typically added to polymer powders for selective laser sintering to increase flowability. These additives reduce friction between particles by electrostatic repulsion, leading to better bed density and part properties. The same repulsion however can hinder particle coalescence in the melt, reducing part density.

The same repulsion however can hinder particle coalescence in the melt, reducing part density. This study investigates the effect of different amounts of flow aid on flowability and coalescence. A polybutylene terephthalate (PBT) powder with spherical morphology, specially designed for selective laser sintering, is used as a base and ideal model material. The coalescence is monitored by hot-stage optical microscopy. It was found that with increasing amounts of flow aid, the flowability could be characterized by three regions; one of low flowability where the powder contained insufficient flow aid, followed by a sharp transition region towards a high flowability region, where the flow aid was most effective. Addition of flow aid impacted the coalescence, which was marked by an increase of the temperature at which the particles started to melt and flow.

Addition of flow aid impacted the coalescence, which was marked by an increase of the temperature at which the particles started to melt and flow. At the maximum concentration of 0.5 wt.-% flow aid, the melting of some particles was delayed, and they remained solid for longer time at temperatures beyond the melting temperature. The optimum amount of flow aid therefore lies in the plateau region of high flowability, but before the occurrence of delayed melting.IntroductionSelective Laser sintering (SLS) is a primary shaping process; the properties of parts depend on many material- and process-based factors. One of these is the flowability of the powder. A key step during the LS process is the deposition of a powder layer, which is typically between 80 and 120 μm thick. The flowability determines to what extent powder can be deposited in a layer, and how densely the powder particles are packed.

The flowability determines to what extent powder can be deposited in a layer, and how densely the powder particles are packed. The density of the powder layer is closely related to the density of produced parts, shown for example by Schmid et al. [1]. To obtain dense parts with optimal mechanical properties, the powder flowability must be optimized. The most widely applied method to improve powder flowability is to add a small amount of nanometer-sized flow aid, demonstrated amongst others in [2–7]. These flow aids are typically a variant of fumed silica and added in quantities between 0.01 % and 0.5 % by weight.Flow aids work based on the principle of electrostatic and steric repulsion, explained in [8]. Without flow aid, polymer particles may cling together because of interparticle van-der- Waals forces, which is elaborated in detail by Rumpf [9]. The presence of flow aid increases the interparticle distance, thereby decreasing the adhesive forces between them.

The presence of flow aid increases the interparticle distance, thereby decreasing the adhesive forces between them. On a macroscopic scale, this improves powder flow. The right choice of flow aid is critically important. Blümel et al. [7] added 0.5 wt.-% of hydrophobic and hydrophilic flow aid to high density polyethylene (HDPE) powder. The powder flowability was measured with a shear cell806and by determination of the bulk density. They found that while the hydrophobic flow aid drastically increased the powder flow properties, the hydrophilic flow aid had no significant, or in the case of the bulk density, even detrimental effect. Most authors opted for hydrophobic flow aids and obtained their desired results. Lexow and Drummer [2] measured an increase of bulk density of a cryogenically ground polypropylene (PP) powder upon addition of 0.1 wt.-% Aerosil® 8200.

Lexow and Drummer [2] measured an increase of bulk density of a cryogenically ground polypropylene (PP) powder upon addition of 0.1 wt.-% Aerosil® 8200. The influence of varying amounts of Aerosil® R106 on the flowability of cryomilled HDPE and PP powders, expressed amongst others by the Hausner ratio, was investigated by Laumer et al. [3]. They found an optimal flowability for HDPE at 0.25 wt.-% flow aid concentration. Addition of more flow aid was less effective. This was attributed to saturation of the particle surface, and subsequent agglomeration of the flow aid itself. The same effect was not observed for the PP powder; a continuous increase of powder flowability up to concentrations of 1.0 wt.-% flow aid was reported. To improve the flowability of a cryogenically ground PBT copolymer powder, Arai et al. [4] added 0.1 wt.-% Aerosil® R200H.

To improve the flowability of a cryogenically ground PBT copolymer powder, Arai et al. [4] added 0.1 wt.-% Aerosil® R200H. They observed enhanced flowability, and were able to determine that the flow aid did not have any effect on the thermal properties of the material.To achieve high density laser sintered parts, not only the flow behavior of the powder is important, but also the rheology and surface tension of the polymer in the melt, which is described by Schmid [10]. Fumed silica’s are also used as additives to influence rheology, and to create materials with thixotropic behavior. Even though the flow aid concentration in SLS materials is relatively low compared to the concentration used to influence rheology, it is imaginable that particle coalescence in the melt can be affected by the flow aid presence. Benedetti et al. summarized the usability of various evaluation methods based on different optical features of particles in hot-stage optical microscopy (HS-OM) [11].

Benedetti et al. summarized the usability of various evaluation methods based on different optical features of particles in hot-stage optical microscopy (HS-OM) [11]. The coalescence of polymer SLS powders has been investigated by various other researchers. The coalescence rate of three polyether(ether)ketone (PE(E)K) powders with different viscosities was determined by Berretta et al. [12], by tracking the neck formation with HS-OM. Particle coalescence was slower for the grades with higher viscosities, but the interpretation of the results is difficult due to the irregular shape of the cryomilled particles. Verbelen et al. [13] tracked the coalescence of different commercial polyamide-based SLS powders with HS-OM. In a later study, Dadbakhsh et al. [14] used HS-OM to view the coalescence of polyamide-12 powders at different degrees of aging.

In a later study, Dadbakhsh et al. [14] used HS-OM to view the coalescence of polyamide-12 powders at different degrees of aging. They identified a slow softening region up to the onset melting temperature of the polymer, followed by a fast coalescence and a subsequent molten phase region. Coalescence occurred slower for powders that had aged more, which was attributed to the increased molecular weight and associated melt viscosity due to the aging. Hejmady et al. [15] used a special setup to observe the coalescence of two laser sintering particles as they are molten by laser irradiation, providing further insight to the melting and coalescence behavior of particles during the actual laser sintering process.HS-OM has been identified as a method to both qualitatively and quantitatively assess the coalescence of polymer powders. It is used in this work to assess the effect of different amounts of flow aid on particle coalescence of a spherical polybutylene terephthalate (PBT) powder.

It is used in this work to assess the effect of different amounts of flow aid on particle coalescence of a spherical polybutylene terephthalate (PBT) powder. Complemented with determination of the powder flowability as expressed by the Hausner ratio, as well as the avalanche angle and avalanche fractal from powder rotational measurements, the optimal concentration can be determined. Comparable results are, according to the authors’ knowledge, not available in present literature.807Materials and MethodsMaterials. PBT TORAYCON 1200M (TORAY, Tokyo, Japan) was obtained in granulate form. The material did not contain any stabilizers or other additives to prevent these from affecting the SLS process. Hydrophobic fumed silica Aerosil® R812 flow aid (Evonik Ind., Essen, Germany) was obtained and used without further modification.

Hydrophobic fumed silica Aerosil® R812 flow aid (Evonik Ind., Essen, Germany) was obtained and used without further modification. Aerosil R812 is a trimethyl terminated fumed silica, and considered a good general choice for improving the flowability of different polymer powders.Powder Production. The PBT powder was produced via melt emulsification. During this process, the target material is melt blended in a single-screw extruder with a water-soluble resin. The production of the particular powder used in this study is extensively described in [16] Material preparation. The PBT powder was dry blended with R812 flow aid in a rotating mixing drum with a diameter of 200 mm. The drum shown in figures 1 and 2, was specially designed to facilitate mixing of two components. Each blend was mixed for 30 minutes with a rotation speed of 60 rpm. First, a masterbatch containing 0.5 wt.-% of flow aid was prepared. This masterbatch was used to prepare the blends listed in table 1.

First, a masterbatch containing 0.5 wt.-% of flow aid was prepared. This masterbatch was used to prepare the blends listed in table 1. The weight of each blend totaled 100 g.Table 1. Prepared blends of PBT and flow aid.DesignationPBT_0.5 PBT_0.1 PBT_0.05 PBT_0.01 PBT_0.005 PBT_VirginFlow aid concentration [wt.-%] 0.5 0.1 0.05 0.01 0.005 0.0CommentMasterbatchNo flow aidFigure 1. Schematic cross-section of the used mixing drumFigure 2. Mixing drum on mixing stage (shown here without lid)Particle size distribution. The particle size distribution was measured by dynamic light scattering on an LS230 (Beckman Coulter, Brea, CA, United States) instrument. Approximately 0.1 g of powder was added to approximately 20 ml acetone and thoroughly stirred. The particle size was calculated based on the Fraunhofer diffraction model, for the range between 0.4–2000 μm.Scanning electron microscopy.

The particle size was calculated based on the Fraunhofer diffraction model, for the range between 0.4–2000 μm.Scanning electron microscopy. Scanning electron microscopy images were obtained on a JSM 7100F scanning electron microscope (JEOL, Tokyo, Japan). Samples were deposited on a piece of carbon tape on an aluminium SEM stub and sputter coated with a 10 nm layer of Pt/Pd (80/20).808Thermal analysis. Differential Scanning Calorimetry (DSC) measurements were carried out on a DSC 25 (TA Instruments, New Castle, DE, United States). All measurements were conducted under a nitrogen atmosphere, from 25°C to 250°C, with heating and cooling rates of 10°C/min.Powder flowability. The powder flowability as expressed by the Hausner ratio was determined by measurement of the powder bulk and tapped density. The bulk density was determined by filling a 25 ml brass cup with powder through a Carney funnel and recording the weight. Each powder blend was measured three times.

The bulk density was determined by filling a 25 ml brass cup with powder through a Carney funnel and recording the weight. Each powder blend was measured three times. The tapped density was determined with a BeDensiT3 Tap Density Volumeter (Dandong Bettersize Instruments, Liaoning, China). One measurement consisted of filling three measuring cylinders with 25 g of powder each. The cylinders were then automatically tapped 1’500 times with a frequency of two taps per second. The powder volume at the end of the tapping cycle divided by its weight gives the powder tapped density.The powder flowability was also evaluated with a Revolution Powder Analyzer (RPA) (Mercury Scientific, Newtown, CT, United States). An exact measure of 25 mL, respectively 21.8 g powder at tapped density was added to a drum with a diameter of 50 mm. The drum was rotated with a speed of 0.6 rpm, while a camera recorded 384 avalanche events.

The drum was rotated with a speed of 0.6 rpm, while a camera recorded 384 avalanche events. Each time directly following an avalanche, the avalanche angle and roughness of the powder surface were evaluated.Optical microscopy. Optical microscopy was performed with a DM6 optical microscope (Leica Microsystems, Wetzlar, Germany) in transmitted light mode with 100 times magnification. For the evaluation of coalescence behavior, a THMS 600 hot stage (Linkam Scientific Instruments, Tadworth, United Kingdom) was installed. The following temperature program was used: (1) 25 – 210°C @ 50°C/min; (2) isothermal 5 min; (3) 210 – 230°C/min @ 5°C/min; (4) isothermal 10 min. An image was acquired every four seconds.Results and DiscussionParticle size and shape The volume-based size distribution of the starting virgin powder, measured with dynamic light scattering, is shown in figure 3. A typical SLS powder exhibits a particle size distribution that stretches from 10 to 150 μm [10].

A typical SLS powder exhibits a particle size distribution that stretches from 10 to 150 μm [10]. In case of the PBT powder under investigation, 90% of the particle size by volume falls within this range. The powder has a D10 of 10.0 μm, a D50 of 41.5 μm, and a D90 of 100.8 μm.- Differ ntial distributi~n - - C umu ative distributi n -D10_v =10 . µm D50.v =41 . µm 0 90_V = 100 8 µm(I) E :::,~ :i:: i:5 2(I) E :::, 50 o > E :::, uo -1--~=..1---.,...;:c,..;....1---.;:::.:;:..,.,=.1--~~ -.1--....-->~=1--~ ---,-1- 0 0.010.1100100010Particle Diameter [µm]Figure 3. Differential (black, solid line) and cumulative (blue, dashed line) size distribution of the starting PBT powder.809The particle shape and morphology was qualitatively evaluated with SEM, two images are shown in figures 4 and 5. The particles have a clear spherical shape. The magnified image in figure 5 shows that the surface of the particles is smooth, no porosity neither satellites are observed.

The particles have a clear spherical shape. The magnified image in figure 5 shows that the surface of the particles is smooth, no porosity neither satellites are observed. As is also indicated by the small initial peak around 1 μm in the size distribution in figure 3, there is a considerable amount of small particles, which mainly seem to be clustered around the larger particles.Figure 4. Overview of spherical PBT particlesFigure 5. Detail of PBT particlesThermal properties The first heating and cooling DSC traces of the investigated PBT and flow aid blends are shown in figure 6. The curves have been overlaid to emphasize any differences. The melting peaks in the first heating curves overlap almost perfectly, indicating that the addition of flow aid does not have any effect on the macroscopic melting behavior. Similarly, no significant effect of the presence of flow aid on the crystallization during cooling is observed.

Similarly, no significant effect of the presence of flow aid on the crystallization during cooling is observed. These results correspond well with findings from Arai et al. [4], who also did not find any difference in thermal behavior between PBT powders with and without 0.1 wt.-% Aerosil®RA200H flow aid.CoolingI~ / I \.I.PST_ V1fg1n - PBT_0.005 - PBT_0f1 - PBT_0. 5 - t-'B,_ u. - PBT_0.1st Heating I1I"' ~~\)2180185190195200205210215220225230Temperature i ("C )Figure 6. 1st Heating and cooling DSC traces of PBT with different amounts of flow aid.810235Powder FlowabilityThe bulk, free-flowing density of the powder as a function of the amount of flow aid is shown in figure 7. The virgin powder with no flow aid has a low bulk density, indicative of a poorly flowing powder. Addition of 0.005 and 0.01 wt.-% flow aid does not contribute much to increasing density.

- - - - :, l1l I\- - - -+ - -\\ •,_1.2.,\ - ·· ·1.1 1- +-------------,--....... - - , - - - -0.01E-30010 10.01 E-30.010 .1Flow Aid Concentration [wt.-%]Flow Aid Co ncentration [wt.-%]Figure 7. Powder bulk, free-flowing density for powders with different flow aid concentration. Dashed lines between data points are to guide the eyeFigure 8. Hausner Ratio for powders with different flow aid concentration. Dashed lines between data points are to guide the eyeSimilar behavior can be observed in the development of the Hausner ratio with different amounts of flow aid, shown in figure 8. Two plateaus of low and high flowability exist, with a narrow transition region in between. The Hausner ratio shows the same behavior as the bulk density as it combines the bulk density with the tapped density.

The Hausner ratio shows the same behavior as the bulk density as it combines the bulk density with the tapped density. Even though the powder with 0.05 wt.-% or more flow aid packs more densely in its free-flowing state than powder with no or low amounts of flow aid, the size and shape of the particles is the same for all powders. The tapped density depends on these two factors. After a sufficient amount of taps, the tapped density for a powder with the same size and shape of particles, is the same.(a) PBT_Virgin(b) PBT_0.005(c) PBT_0.01(d) PBT_0.05(e) PBT_0.1(f) PBT_0.5Figure 9. Snapshots of powders in the Revolution Powder Analyzer (RPA) drum.The powder flowability was also investigated in a Revolution Powder Analyzer (RPA). The RPA consists of a rotating drum with transparent lids, through which the powder flow is evaluated with a digital camera and image recognition software.

The RPA consists of a rotating drum with transparent lids, through which the powder flow is evaluated with a digital camera and image recognition software. Amongst other parameters, the RPA allows for the determination of the powder avalanche angle (cid:302)A and the avalanche fractal fA. The avalanche angle represents the resistance of the powder to flow, which strongly depends on the interparticle friction. The smoothness of the powder surface after an avalanche is expressed by the surface fractal, which is not only governed by the interparticle fraction, but also dependent on the particle size and geometry. Detailed investigations regarding powder811characteristics and respective flow properties were carried out by Amado et al. [17] and Vetterli [18].Figure 9 shows snapshots of the powder as seen by the camera through the drum during a measurement.

[17] and Vetterli [18].Figure 9 shows snapshots of the powder as seen by the camera through the drum during a measurement. With increasing flow aid amount, the powder surface transitions from very irregular for blends containing 0.01 wt.-% and less flow aid (figure 9a-c), to smooth for blends containing more than 0.01 wt.-% (figure 9d-f). The powder also sticks to the surface all around the sides of the drum for blends with equal or less than 0.01 wt.-% flow aid. Visually, it is already possible to distinguish between two flow regimes; one with poor powder flowability where the flow aid concentration is 0.01 wt.-% or less, and one with better powder flowability at higher amounts of flow aid.655 -'--I601....."" 0) 55 C: <: "' " 50 C: ~ ro > <:.s::45.. - --- ··--·-- ··•-- · - ·•····· - ··•--. ■..40350 .011·~~-~~+--~~-~.,.,--~---~0.010.111-~~-~------~~--~-~0.00.010.1Flow Aid Concentration [wt.-%]Flow Aid Concentration [wt.-%]Figure 10.

■..40350 .011·~~-~~+--~~-~.,.,--~---~0.010.111-~~-~------~~--~-~0.00.010.1Flow Aid Concentration [wt.-%]Flow Aid Concentration [wt.-%]Figure 10. Average avalanche angle for powders with different flow aid concentration. Dashed lines between data points are to guide the eye. The shaded region indicates the width of the avalanche angle distribution as expressed by its standard deviation.Figure 11. Average surface fractal for powders with different flow aid concentration. Dashed lines between data points are to guide the eye. The shaded region indicates the width of the surface fractal distribution as expressed by its standard deviation.Both the avalanche angle and surface fractal, shown in figures 10 and 11 respectively, show a similar trend in dependence of flow aid concentration as the Hausner ratio.

In addition to the presence of two plateaus of low and high flowability with a transition region in between, the consistency of powder flow gradually increases with increasing flow aid concentration. This is indicated by the shaded region in figures 10 and 11. The region width is equal to the standard deviation of the avalanche angle and fractal distributions. A powder flows better and more consistently when these distributions are narrow. The high avalanche angle for PBT_0.01 shows essentially the same cohesiveness, initial resistance to powder flow as PBT_Virgin without flow aid. The associated avalanche fractal however indicates that the powder surface is smoother for PBT_0.01, showing that the flow aid is mainly acting when the powder is in movement.

The associated avalanche fractal however indicates that the powder surface is smoother for PBT_0.01, showing that the flow aid is mainly acting when the powder is in movement. The gradual decrease of the surface fractal standard deviation with increasing amounts of flow aid very neatly describes its effect on the powder flowability.The existence of the two flowability plateaus and transition region can be explained by the working principle of flow aids. As discussed in the introduction, flow aids reduce friction between particles by electrostatic repulsion. There is an optimal flow aid concentration, close to the point where the particle surface is saturated. Less flow aid means the particles still rub against their own surfaces during flow, increasing friction and decreasing bulk density. More flow aid is not contributing to better flowability, because the particle surface is already saturated.

More flow aid is not contributing to better flowability, because the particle surface is already saturated. This behavior is very similar to what happens with the surface tension of a liquid when a surfactant is added, until a certain critical micelle concentration is reached, described in [8,19].812Particle Coalescence The coalescence of powder particles was tracked with HS-OM. The first column of figure 12 shows micrographs of the particles at 210°C, at the start of the measurement. Micrographs of the powders at increasing temperatures are shown in the subsequent columns. Compared to all other powders that contain flow aid, the particles of the virgin blend start to melt and coalesce at the lowest temperature..: , .. 0 • . ·-~ .. ... :fj:t: ... •08..0 PO 'O ' ·~ I.00n i g r ie~00IV _ T B Pe00,/0\0210°C224°C226°C230°C5 0 0 . 0 _ T B P1 0 . 0 _ T B P.. . . .. ~~ ... • ..... " JU~: • . .. . •·· .-.~--- • •• • ~-•.::· ...

At this temperature, coalescence has also progressed considerably for the PBT_0.005 powder, yet some non-molten particles remain. At increasing amounts of flow aid, smaller amounts of powder are flowing and coalescing.The first melting curves of the DSC measurements show no effect of the flow aid on the onset of melting. This however is in contrast to the observations in HS-OM. The powders should all melt at the same temperature, but they do not seem to do so. There are two possible explanations for this effect. One reason could be differences in heat transfer to the particles. The particles are laying on a glass slide, heat is only supplied from the bottom through the slide. Since the particles are spherical, the contact area is quite small. For the smallest particles, this is no limiting factor. As can be observed, these are always the first to melt.

Since the particles are spherical, the contact area is quite small. For the smallest particles, this is no limiting factor. As can be observed, these are always the first to melt. If however a large particle with a much lower surface-to-volume ratio, is laying by itself, untouched by other, smaller particles, it can take a reasonable time before enough heat is transferred to melt it completely. Addition of flow aid separates particles from each other, which may explain why PBT_Virgin and PBT_0.005 melt relatively quickly; their particles are clustered together. Powders with higher amounts of flow aid are more dispersed, and therefore need more time and temperature to melt.One observation that seems to contradict this explanation is the melting behavior of PBT_0.5 in the bottom row of figure 12. At 230°C, a considerable amount of non-molten particles is still present.

At 230°C, a considerable amount of non-molten particles is still present. Some of these are surrounded by polymer melt, which should facilitate heat transfer, yet they melt only after several minutes at 230°C. At this temperature, no gradual melting is observed, but the particles seem to burst and disappear almost instantaneously. In this case, the flow aid can act as a thermal insulator. In sufficient quantities, the flow aid particles form a dense network with a structure comparable to the structure of aerogels. As described in [20], aerogels have an extremely low thermal conductivity. To investigate the amount of flow aid on the surface of particles, SEM images were acquired of PBT_Virgin, PBT_0.05, and PBT_0.5. The images are shown in figure 13.(a) PBT_Virgin(b) PBT_0.05(c) PBT_0.5Figure 13. SEM images at 20’000x magnification of virgin and blended PBT powders.

The images are shown in figure 13.(a) PBT_Virgin(b) PBT_0.05(c) PBT_0.5Figure 13. SEM images at 20’000x magnification of virgin and blended PBT powders. Length of the scale bar is equal to 1 μmAs expected, no flow aid particles are visible on the surface of PBT_Virgin particles. Upon close inspection, flow aid particles are present in PBT_0.05, but they do not saturate the surface. For PBT_0.5 however, the flow aid completely covers the surface of the particles. Additionally, the flow aid is present in the form of aggregates between the particles. This confirms the similarity of the flow aid nanopowder and an aerogel in terms of morphology.The delayed melting is demonstrated by PBT_0.5 only. The other blends melt more gradually, even though that tends to happen at increasing temperatures with increasing flow aid concentration. In terms of powder flowability, at 0.1 and 0.5 wt.-% a plateau is reached, after814which it is unlikely to see drastic improvements.

In terms of powder flowability, at 0.1 and 0.5 wt.-% a plateau is reached, after814which it is unlikely to see drastic improvements. This indicates the particle surface starts to be saturated with flow aid only at these high amounts. PBT_0.1 does not exhibit delayed melting, but does show show optimal flowability in terms of Hausner ratio, avalanche angle, and avalanche fractal. The optimal flow aid concentration likely lies around this point.Conclusions and OutlookThe effect of different amounts of flow aid on the flowability and coalescence of a PBT polymer powder consisting of spherical particles was determined. The flowability as a function of flow aid concentration can be divided into three regions; the first consists of a plateau of low flowability where there is no measurable effect of flow aid addition, up to amounts of approximately 0.01 wt.-%.

This is followed by a sharp transition region over the range between 0.01 and 0.05 wt.-% after which another plateau region of high flowability for concentrations higher than 0.1 wt.-% is reached. Addition of even the smallest amount of flow aid shifts the the coalescence of particles to a higher temperature by approximately 2°C. With increasing amounts of flow aid, coalescence is delayed further. A special case occurs for PBT with 0.5 wt.-% flow aid. Some particles remain solid at high temperatures, disappearing almost instantaneously after a few minutes. Both the flowability and melting behaviors show analogies with surfactant behavior. Mechanisms that hold for surfactants may therefore be applied to systems containing flow aids, which can be helpful to determine the optimal amount of flow aid to add to a powder.The current work summarized the results from initial investigations into powder flow and coalescence behavior as a function of flow aid concentration.

There are multiple possibilities to extend the study to gain a better understanding of the interplay between powders and flow aids. An obvious start is to investigate the effect of flow aids with different chemistry. The Aerosil ® R 812 used in this study is a general-purpose hydrophobic flow aid, but there are many other variants that are likely to have a different influence on powder coalescence and flow. To increase the resolution and validity of the results of the current study, it is necessary to investigate additional blends with different flow aid concentrations. The most interesting blends lie in the flowability transition region between 0.01 and 0.05 wt.Although it is possible to distinguish the differences in flowing behavior of powders containing different amounts of flow aid visually, quantification based on the current results is difficult. The particle size is related to the observed melting temperature.

The particle size is related to the observed melting temperature. That also means that in order to evaluate the results properly, only particles with similar size should be considered. This was not possible with the current measurements, but it is possible to track the coalescence of two similar particles at higher magnification, similar to the method of Berretta et al. [12]. These measurements can also be done at lower heating rates such as 0.5°C/min. Coalescence then occurs quasi-isothermally, and can be related to existing models for the coalescence of two viscoelastic particles. In this way, the dominating effects can be determined and help to explain the observed behavior.Finally, it is important to establish the link between the results measured in the laboratory and the performance of the powder in the SLS process. This was not possible in the current study due to the limited amount of available material.

This was not possible in the current study due to the limited amount of available material. Solutions that allow for the evaluation of part density also for small amounts of powder are currently being developed, and will be used in further research.815AcknowledgementsLuiz Morales of the Scientific Center for Optical and Electron Microscopy (ScopeM) atETH Zürich is acknowledged for his assistance with acquiring SEM images.References1.2.3.4.5.6.7.8. 9. 10. 11.Schmid, M.; Kleijnen, R.; Vetterli, M.; Wegener, K. Influence of the origin of polyamide 12 powder on the laser sintering process and laser sintered parts. Appl. Sci. 2017, 7, DOI: 10.3390/app7050462. Lexow, M.M.; Drummer, D. New Materials for SLS: The Use of Antistatic and Flow Agents. J. Powder Technol. 2016, 2016, 1–9, DOI: 10.1155/2016/4101089. Laumer, T.; Stichel, T.; Raths, M.; Schmidt, M.

E IMPACT OF IN-HOUSE RP UPON FINAL YEAR INDUSTRIAL DESIGN STUDENT PROJECTSR. I. Campbell and A. R. Hodgson Department of Design and Technology, Loughborough University, LE11 3TU, UK Reviewed, accepted August 13, 2003 AbstractA vital part of the Industrial Design and Technology degree course at Loughborough University is the development of appearance and functional prototypes. Previously, the use of rapid prototyping (RP) has been fairly limited because of the high cost and problematic timing issues of using external facilities. However, during 2003, the availability of an in-house FDM 2000 machine has greatly increased the use of RP models. This paper identifies the impact this has had upon a) the design approach, b) the characteristics of the models produced and c) the use of the models with user evaluation trials.

This paper identifies the impact this has had upon a) the design approach, b) the characteristics of the models produced and c) the use of the models with user evaluation trials. Conclusions drawn show that RP has brought major benefits but that there are some pitfalls that must be avoided.IntroductionThe Industrial Design and Technology degree course at Loughborough University is characterised by an integrated treatment of the aesthetic, ergonomic and functional aspects of product design. Final year students are encouraged to incorporate all three into their major design project and to validate their product design through user evaluation. This necessitates the use of appearance models and functional prototypes within user evaluation trials. Conventionally, appearance models will be used to convey aesthetics and some ergonomic aspects to the user while functional prototypes are used to convey the remaining, more technical features.

Conventionally, appearance models will be used to convey aesthetics and some ergonomic aspects to the user while functional prototypes are used to convey the remaining, more technical features. In industry, rapid prototyping (RP) has been used to help produce both appearance models and functional prototypes, often yielding significant time and cost savings. However, recent research at Loughborough [Evans, 2003] has shown that the greatest benefit of RP to industrial design is the ability to create an integrated appearance model/functional prototype, often called an appearance prototype. When used in this manner, it can not only reduce time and costs but also improve the designer's confidence in the results of user trials because of the increased realism of the prototypes used.

When used in this manner, it can not only reduce time and costs but also improve the designer's confidence in the results of user trials because of the increased realism of the prototypes used. Therefore, it has become an aim of the Department of Design and Technology (DDT) at Loughborough to encourage the use of RP whenever possible and beneficial.If rapid prototyping (RP) is to become a widely used tool within the field of industrial design, it is necessary for students of the subject to be completely familiar with its capabilities and limitations. To this end, DDT set up close relationships with local RP service providers, initially in a nearby RP bureau, and more recently, in another department at Loughborough. However, the uptake of RP remained rather low, with less than 10% of final year students making use of it during their major design project.

However, the uptake of RP remained rather low, with less than 10% of final year students making use of it during their major design project. This was partly due to the relatively high cost of using the RP bureau and some scheduling problems with the other university department. However, it was also believed that there may have been a lack of willingness to use RP techniques because they were seen as "engineering only" tools. At the beginning of the 2002/03 academic year, DDT acquired FDM 2000 and ThermoJet RP machines for its own use. Although quite old, the FDM126machine worked well, once a new deposition head had been fitted. It was hoped that the addition of an in-house facility would promote greater student familiarity with RP techniques in general and break down any perceived barriers. This paper describes the effect that in-house RP has had within DDT and the impact it has had upon the students' design process.

This paper describes the effect that in-house RP has had within DDT and the impact it has had upon the students' design process. The overall aim of the paper is to present results and draw conclusions that will help industrial design departments in other universities who find themselves in a similar situation.BackgroundThe Industrial Design and Technology course at Loughborough is a three year undergraduate programme that can lead to a BSc or BA, depending on which optional modules are taken. It is one of only a few university courses in the UK which place major emphasis on both the outward, aesthetic properties of a product together with the internal technology required to provide the product's function. Therefore, the students cover a wide range of subjects including sketching, manual model making, product semantics, multi-media presentation, engineering drawing, electronics, mechanics, materials, computer-aided design (CAD) and design for manufacture.

Great importance is attached to a hands-on approach to design. Students undertake project work in all three years aimed at assimilating and applying the knowledge they have gained from other modules. Graduates from the programme have been known to embarque on careers in engineering design as well as industrial design.The major design project undertaken by all final year students is the culmination of all their practical activities. It runs across both semesters and has a nominal student time input of 500 hours, in total. The first semester concentrates on finding a potential market for a product, developing a product design specification, generating several concept designs and selecting one for further development. During semester two the students are expected to convert the chosen concept design into a fully detailed design, produce a functional prototype, have it evaluated by potential users and consider design for manufacture issues.

If the functional prototype does not reflect the actual appearance of the final product then a student will be expected to produce an It is during this "prototype manufacture" stage of the project that appearance model also. students will employ their model making skills. Traditionally, this has involved manual modelling in blue foam, medium density fibreboard (MDF), various plastics, aluminium, steel and occasionally, textiles. The techniques used include machining, forming, welding, filling, sanding and painting. Computer numerical control (CNC) milling and routing is sometimes used for both component and tool manufacture, and more recently, vacuum casting has been used to replicate models in more representative materials.

In the past, the predominant use of manual techniques has often made it difficult for students to combine both an organic form and intricate internal details within a single appearance prototype.Throughout the duration of the project, students are expected to attend weekly seminars that cover both organisational aspects of the module and various skills and knowledge they will need to apply. RP is introduced to the students during one of these seminars. They are given an overview of how it works, the techniques available, its pros and cons, and the importance of having a high quality CAD model to drive the RP process is stressed. All of the students will have covered 3D CAD in both their first and second year but mostly this is confined to solid modelling techniques. It should be noted that around half of the students will already have had a127more detailed introduction to RP during an optional final year module entitled "Computer-aided Modelling and Manufacture" (CAMM).

This module also exposes them to more advanced CAD/CAM techniques such as freeform surface modelling and three axis CNC machining. Therefore, the CAMM students tend to have acquired more skills and knowledge in regard to the "digital" design process. It is interesting to note that 85% of the students who used RP during their project had attended the CAMM module.Data CollectionThe aim of this research was to determine the effect that the introduction of in-house RP has had within DDT and, in particular, the impact it has had upon the students' design process. To do this it was necessary to gain an insight into how and why students had used RP. This was undertaken in two ways. Firstly, a questionnaire was given to all the final year students asking specific questions about their use (or non-use) of RP. The aim of the questionnaire was to obtain some quantitative data regarding the use of RP.

The aim of the questionnaire was to obtain some quantitative data regarding the use of RP. The questionnaire generally followed the same format as one that had been used with industry-based designers in a previous research project [Campbell, 1998]. However, it had to be tailored to the different circumstances faced by students. The questions are listed in the Appendix. Secondly, to gain more qualitative data on the impact that RP had had upon students' design process, several student projects were treated as case studies. Their use of RP was tracked throughout their projects and the resulting prototypes analysed to see what benefits RP had yielded. The results obtained from both strands of the research are now presented together with some preliminary analysis.Questionnaire Results and AnalysisThe questionnaire was issued to all 90 final year Industrial Design students who had just completed their major individual design project.

35 students completed the questionnaire, equivalent to a response rate of 39%. Of these 35 students, 20 had used RP during their project whereas 15 had not. It is known that a large majority of those who did not respond had not used RP. Therefore, it can be estimated that approximately 25% of the final year had used RP this year, a big increase over the previous year when only around 10% had used it. A range of reasons where given for not using RP with just under half of the responses being “not appropriate to project”. This is a fairly vague phrase but other students were more specific. Virtually equal numbers cited “lack of know-how”, “unsuitable materials”, “product too large”, “RP too expensive” and “simpler to use conventional techniques” as reasons for not using RP. In regard to RP being too expensive, the students were charged at a rate of 25 pence (40 US cents) per gramme for FDM models.

In regard to RP being too expensive, the students were charged at a rate of 25 pence (40 US cents) per gramme for FDM models. This was much less than a commercial rate but could still result in larger models having a significant cost. The “lack of know-how” most likely refers to a deficiency in CAD skills rather than a lack of knowledge about RP techniques. This can be assumed because only half of the students took the optional final year CAMM module.Of the 20 students who did use RP, the predominant reason given (cited by 14 students) was “the ability to create more complex or more organic product forms”. Other reasons were “improved accuracy”, “better surface finish”, “faster process”, “more convenient” and “to gain experience”. None of these was cited by more than four students.

Other reasons were “improved accuracy”, “better surface finish”, “faster process”, “more convenient” and “to gain experience”. None of these was cited by more than four students. It should be understood that128the majority of the students were comparing RP to conventional, manual techniques for model making rather than to CNC machining or other final production processes.Half of the students required only a small number of RP models (five or less) and the distribution for the remainder can be seen in Figure 1. The largest number of models made for any one student was 26. 11 of the students made use of the in-house FDM machine, whereas the numbers who used SLA, ThermoJet and SLS were 8, 4 and 2 respectively (7 students used more than one machine). Although the cost of using FDM was greater (SLA, ThermoJet and SLS models were provided for free), the convenience of using locally located facilities seemed to be more important to many.

Although the cost of using FDM was greater (SLA, ThermoJet and SLS models were provided for free), the convenience of using locally located facilities seemed to be more important to many. Other reasons given for usage of particular techniques were accuracy, strength, surface finish and suitability for investment casting. Indeed the fact that several students used more than one technique showed that they had gained an understanding of the relative merits of the various systems available.s t n e d u t s12108f o r e b m u N64201 to 55 to 1011 to 1516+Number of RP models usedFigure 1: Number of students against numbers of models used.75% of the students used their RP models as functional components within their prototype whereas only 20% used models for purely visualisation purposes. This shows that the majority of students had recognised the ability of RP to create models that have similar characteristics to final production components.

This shows that the majority of students had recognised the ability of RP to create models that have similar characteristics to final production components. 25% of students used their RP models for vacuum casting, again the aim being to arrive at functional components. Other uses for the models where investment casting (1 student), assembly checking (2 students) and ergonomic testing (1 student). Several students used different models for different purposes, again showing a good understanding of RP system capabilities.In an attempt to determine if RP had opened up new design possibilities for the students, they were asked if they would still have made the same models if RP had not been available. Unfortunately, nine of students answered "not sure" which does not give much indication of their thought processes. However, four students said they would have made do without a physical model, perhaps using 2D images or a CAD animation instead of an appearance model.

However, four students said they would have made do without a physical model, perhaps using 2D images or a CAD animation instead of an appearance model. Obviously, for these students, RP had enabled them to produce a more tangible output from their design process. For the seven students who said they would definitely have made the same models, all of them said they would have used manual modelling techniques with three saying they would also have used CNC machining. Following on from this, the students were asked if they had used conventional modelling techniques alongside RP. 85% of them had done so and129the number of students using each of a wide range of techniques is shown in Figure 2. This shows that the traditional skills of industrial designers compliment well the capabilities of RP. The students were entirely at ease with creating a high quality prototype from an RP model with stair-stepping and supports.

The students were entirely at ease with creating a high quality prototype from an RP model with stair-stepping and supports. Furthermore, the technological capabilities of the students enabled them to combine this model with others created using plastic and metal-forming techniques.s t n e d u t sf o r e b m u N10 8 6 4 2 0FillingS andingS prayingV ac formingV ac castingW eldingM achiningC N CFigure 2: Conventional modelling techniques used alongside RPThe students reported a variable degree of success with using RP. The proportion of RP models that were deemed “satisfactory” by each student varied from 100% to zero. Figure 3 shows the wide distribution of satisfaction levels. Most students have had a reasonably favourable experience in using RP but one student was dissatisfied with all the models produced. This was partly due to different expectations from the models and partly from the different RP techniques used.

This was partly due to different expectations from the models and partly from the different RP techniques used. 70% of the students judged the quality of their RP models by the level of accuracy, 50% by appearance and 40% by surface finish (Most students listed more than one criteria). However, the strength, functionality and surface texture of the RP models were also considered important by some students. RP models produced by some techniques do not measure up well against all of these criteria. The two predominant reasons given for lack of satisfaction were “poor surface finish” (cited by eight students) and “poor accuracy” (cited by seven students.

The two predominant reasons given for lack of satisfaction were “poor surface finish” (cited by eight students) and “poor accuracy” (cited by seven students. Considering that some students had used RP specifically to improve upon these two factors, it is possible they held some rather unrealistic expectations that have not been met.s t n e d u t sf o r e b m u N9 8 7 6 5 4 3 2 1 00 to 20% 21 to 40% 41 to 60% 61 to 80% 81to 100%Figure 3: Proportion of RP models judged satisfactory by students.130The last question tried to gauge what impact RP had had upon the design process employed by the students. They were asked how RP had changed their approach to design. 25% of students said that RP had enabled them to produce a more realistic prototype. A further 10% stated that they had been able to produce a single appearance prototype.

25% of students said that RP had enabled them to produce a more realistic prototype. A further 10% stated that they had been able to produce a single appearance prototype. When taken with the 15% who cited “more complex shape possible” the result is that half of the students had been able to change their physical modelling strategy because of RP. The only other change in design process that was cited by more than one student was that model making had been quicker. This may also refer to the fact that only one set of models was required rather than both an appearance model and a functional prototype.The results of the questionnaire demonstrated that RP had been used intelligently by many students and had proved beneficial to their design project. However, little indication was given as to the extent that RP can change the way students work.

However, little indication was given as to the extent that RP can change the way students work. To determine this, three of the highest ranked student projects were taken as case studies and analysed in more detail.Case StudiesThree student projects used as case studies to probe further into issues that had been raised by the questionnaire. These were a) changes in the design approach followed, b) changes in the characteristics of the models produced and c) the enhanced use of the models within user evaluation trials. Each of the three case studies is now discussed in turn.Case study 1 – Michael’s Glue GunMichael's product was a re-design of a glue gun that was aimed at the ‘hobbyist’ and ‘craftsman’ who wanted more sophisticated control of the glue as it is dispensed.

His final design integrated mechanical, electronic and visual elements in order to meet his intended design specification (see Figure 4).Figure 4: Final design of glue gun131Early design development raised major problems concerning his aim to incorporate a motorised glue delivery system into a small pen-like product. Instead, it was decided that a separate motor housing should be coupled to the pen via a drive cable, since the required purpose-built motors and mechanical components were outside the scope of his project work. He also decided that his product would be evaluated using two different prototypes.

He also decided that his product would be evaluated using two different prototypes. Firstly, an appearance model that would represent the aesthetics and "feel" of the final design, and secondly, a functional prototype that would represent the functional operation of the product – a key part of the ‘added value’ for this new product.Towards the later stages of the design project, Michael developed two parallel designs using CAD modelling to create a detailed assembly of parts. One design was to represent the new glue gun as it would be commercially manufactured, with custom designed electrical and mechanical parts. The model for this design was to be made using SLA parts. The other design included a similar (though slightly different) dispenser, a simple drive unit that would house a prototype motor drive (from a cordless screwdriver) and prototype electronics circuit.

His intention was to deliver the glue with varying speed, to sense when the user has finished and retract the glue stick to prevent it dripping, and also to vary the heating element power depending on the intended gluing operation. The system required significant development and so a large drive unit, with substantial space for further development, was built in the stronger FDM 2000 material. Figure 5 shows the two different prototypes side by side, and their differences are quite apparent.Figure 5: Appearance model (on left) and functional prototype (on right).Michael made no attempt to finish the surface of the FDM parts used to make the functional prototype. However, the SLA parts for the appearance model of the final design were hand finished and painted. Although not as robust as the FDM model, the purpose of this prototype served only to illustrate form and appearance, and it complemented the functional prototype well.

Although not as robust as the FDM model, the purpose of this prototype served only to illustrate form and appearance, and it complemented the functional prototype well. The main conclusion to be drawn here was:that RP processes can be carefully selected to produce different prototype outcomes.132Case study 2 - Jon's CD PlayerJon used a variety of different RP processes at different phases of his design development work, all linked to his significant CAD modelling capability with Pro/ENGINEER software. He designed a domestic CD player, which included speakers that could be separately located, and also allowed remote control of the CD player volume and track (see Figure 6). The product comprised a main body (that included the CD player itself and the speaker battery chargers) and two symmetrical speaker ‘pods’ (that included audio receivers and control transmitters, together with batteries and NXT sound exciters so that the speaker cases acted as sound boards).

This allowed the CD player to be used in wet environments, like showers, since the speaker ‘pods’ could be completely sealed.Speaker PodMain BodySpineBaseFigure 6: CD Player with various features indicated.Jon made use of ThermoJet wax models for proving the fit of the pods to the main body, and went through a number of design iterations to prove the initial concepts as the shape and form of the product developed. The product functionality took longer than anticipated to complete, so instead of using SLA models and vacuum casting to produce prototype polymer shells, he chose to use FDM models and then finish/paint their surfaces. He considered that the extra time needed to finish the FDM surface would be less than the time required to make silicone moulds and cast resin parts.The two plastic mouldings that formed the main body were attached to a cast aluminium ‘spine’.

Although Jon considered modelling the spine by hand (which would have sufficed for such an early prototype), he decided to use a ThermoJet wax model of the part to make an investment casting in aluminium alloy, made as a one-off part. This provided a rigid framework133onto which the less substantial RP parts could be fastened, and so the CD player could be effectively evaluated as a fully working, appearance prototype.that the faster, less functional RP techniques enable rapid and copious design iterations, • that the need for separate appearance model and functional prototype can be eliminated via the combined use of RP and secondary processes.Case study 3 – Kelly’s Food Preparation DeviceLike Jon, Kelly made use of a several RP techniques at different phases of her design project. She was developing a food preparation product that exploited a new textured plastic within what was aimed to be a “cooking is fun” context.

She was developing a food preparation product that exploited a new textured plastic within what was aimed to be a “cooking is fun” context. As with all students of the Industrial Design and Technology course, she started by analysing the potential market and considering the needs of users, helping her to generate initial ideas, concepts and design specifications. During this time, some mechanical parts were modelled in CAD, and evaluated through mechanical simulation, but she still felt the need to physically model the mechanism.The shape and form of the different parts were developed using a combination of sketch work and CAD modelling, and then built in wax using the ThermoJet process (an example part is shown in Figure 7). These helped Kelly to evaluate the overall appearance of her design together with the comfort and other related ergonomic factors of the hand-held parts in particular. These parts could have been modelled in foam, either by using 3-axis CNC routing, or by hand.

These parts could have been modelled in foam, either by using 3-axis CNC routing, or by hand. However, the wax RP model was an accurately formed prototype part that very closely followed It required minimal finishing and was quickly the CAD model and Kelly’s intended design. achieved using an STL file generated from Pro/ENGINEER. The ability of the model to build internal features in the casing parts was particularly beneficial, since it allowed Kelly to check for fit and location of mechanical parts. Several versions of wax-modelled prototypes were produced during the concept development phase of the project, alongside continuing sketch and CAD modelling work.

Several versions of wax-modelled prototypes were produced during the concept development phase of the project, alongside continuing sketch and CAD modelling work. It is thought that the use of quick and easy RP output enabled a more effective design to be developed through a number of early design iterations.Figure 7: One of the ThermoJet models used for aesthetic, ergonomic and assembly evaluation.134As the project progressed, more robust components became necessary for full assembly and testing purposes and so FDM was used. Kelly decided to hand-finish these parts to in order to present her progress to her external "client company" (see Figure 8).Figure 8: Hand-finished FDM partsHowever, the quality of surface finish and detail was not considered to be sufficient for a final appearance prototype. Following some minor re-design after evaluation of the FDM models, Kelly used SLA models with subsequent vacuum casting for this purpose.

Following some minor re-design after evaluation of the FDM models, Kelly used SLA models with subsequent vacuum casting for this purpose. Figure 9a shows a combination of SLA and vacuum casting parts prior to final finishing.(a)(b)Figure 9: Appearance prototype before and after finishing.The SLA models required less finishing than FDM would have needed, and were vacuum cast in an ABS-replicating resin with appropriate colouring pigment to minimise finishing requirements. This resulted in an appearance prototype that was robust enough to be used by a135number of people in a typical food preparation context (see Figure 9b). The resulting evaluation provided very effective feedback on the user acceptance and operation of the product.

The resulting evaluation provided very effective feedback on the user acceptance and operation of the product. The main conclusions to be drawn from this project were:that the use of RP at different stages in the design process enabled more design iteration and resulted in a higher quality product,that the combination of RP with conventional modelling techniques provided a high quality appearance prototype that was suitable for extensive user trials in a realistic environment.ConclusionsThe availability of an in-house FDM machine served to raise the general profile of RP within the department. As a result, the student use of other RP techniques increased over previous years. Aligned to the increased availability of RP modelling was the development of increased CAD modelling capability. Students were able to capture their design intent in Pro/ENGINEER through a combination of solid modelling, surface modelling and freeform surface modelling (using ISDX).

Students were able to capture their design intent in Pro/ENGINEER through a combination of solid modelling, surface modelling and freeform surface modelling (using ISDX). This was a pre-requisite for their RP work, and vital to the successful implementation of RP in both appearance models and functional prototypes.The questionnaire captured data from most of the students who had used RP for their design project. RP was primarily adopted to achieve more complex and more organic forms that would otherwise be difficult and time consuming. This was only beneficial because the traditional model making skills of industrial designers were used to produce high quality model finishes. Different RP systems were used for different outputs and it was highly desirable that a range of complimentary techniques were available. Some students had unrealistic expectations of what RP could achieve. It is therefore essential that students are well informed about various system capabilities.

Some students had unrealistic expectations of what RP could achieve. It is therefore essential that students are well informed about various system capabilities. Most significantly, half of the students stated that RP had changed their strategy for creating physical models.These results were supported by the case studies, which showed that RP processes had been carefully selected to achieve different outcomes. They also showed that the rapid and copious design iterations facilitated by RP result in more optimised designs. The most important change in modelling strategy was the combination of aesthetics and functionality within a single prototype. This led to more realistic user trials, which in turn, should give more reliable feedback on the product design.ReferencesEvans, M.A. and Campbell, R.I.

This led to more realistic user trials, which in turn, should give more reliable feedback on the product design.ReferencesEvans, M.A. and Campbell, R.I. (2003), "A Comparative Evaluation of Industrial Design Models Produced Using Rapid Prototyping and Workshop-Based Fabrication Techniques", accepted for publication in Rapid Prototyping Journal.Campbell, R.I. (1998), "Using Feature-based Product Modelling to Integrate Design and Rapid Prototyping ", PhD thesis, Manufacturing Engineering and Operations Management, University of Nottingham, UK.136Appendix: Questions asked in the questionnaire1. Did you use one or more rapid prototyping (RP) models for your design project?(If yes, please go to question 5.)2. Did you consider using RP during your design project?(If no, please go to question 4.)3. Why did you decide not to use RP?Thank you for completing the questionnaire.4. Why did you not consider using RP?Thank you for completing the questionnaire.5.

Why did you decide not to use RP?Thank you for completing the questionnaire.4. Why did you not consider using RP?Thank you for completing the questionnaire.5. Why did you decide to make use of RP models?6. How many RP models did you have made for your design project?7. Which RP system(s) were used and why did you select them?8. What design or prototyping purpose did you use your RP models for?9. Would you still have made these models if RP had not been available?10. If yes to question 9, how would you have made your models?11. Besides using RP, did you still make use of other model-making techniques (e.g. to make other parts or to finish your RP models)?12. If yes to question 11, which technique(s) did you use?13. What criteria did you use when evaluating the quality of your RP models?14. What proportion of the RP models you had built were of satisfactory quality for your purpose?15. What were the reasons for any of your models not being satisfactory?16.

What proportion of the RP models you had built were of satisfactory quality for your purpose?15. What were the reasons for any of your models not being satisfactory?16. How has the availability of RP changed your approach to your design project?Thank you for completing the questionnaire.137.

viewed PaperTHE IMPACT OF MEASUREMENT METHODOLOGY ON THE DIAMETER MEASUREMENT OF SIMPLE ADDITIVELY MANUFACTURED FEATURESM. Praniewicz*, J. C. Fox*, C. Saldana†National Institute of Standards and Technology, Gaithersburg, MD 20899, United States † George W. Woodruff School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332, United StatesAbstractDimensional qualification of additive manufacturing (AM) components is a continuing research problem. Different measurement techniques implemented on the same feature can yield different measurement results. While this can also be true for components made from more traditional manufacturing processes, the deviations between measurement techniques are often increased by an order of magnitude due to greater form and surface texture variations that occur on AM components. Understanding the origins of deviations and comparability of measurement processes is crucial to the measurement of AM components.

Understanding the origins of deviations and comparability of measurement processes is crucial to the measurement of AM components. In this work, identically designed components are produced using a laser powder-bed fusion process. The components are then measured using manual gaging and a coordinate measurement machine. The measurement of diameter is executed using various association criteria. A statistical analysis is performed to determine the comparability between the measurement techniques. Results indicate that the selection of different association criteria can provide statistically significant differences in the measurement result.IntroductionAdditive manufacturing (AM) technologies have progressed from limited applications in research and development to applications in industrial production. Both polymer and metallic AM processes are currently used to produce consumer products.

Both polymer and metallic AM processes are currently used to produce consumer products. However, these are typically in applications where there is low risk associated with component failure. Industries with highly stressed, high cycle, or dimensionally critical components, such as the consumer aerospace or automotive industries, have had a slower adoption rate of AM produced components. This is largely due to challenges in the qualification of component geometry and material properties [1]. The complex geometries afforded through the AM process create significant challenge for component inspection. Thus, the geometrical qualification of AM components has required a new, developing, area of research [2].While the end goal of geometric qualification is providing measurements to document product quality, it requires input from all stages of the component development cycle.

Without clear communication between the design, manufacturing, and inspection phases of the component, there is potential for the unnecessary production of initial components that don’t comply with the component requirements and necessitates an altered design, manufacturing, or inspection to produce a second batch of components. This undermines the capability of additive manufacturing to rapidly produce complex components and adds additional development, labor, and material costs.

This undermines the capability of additive manufacturing to rapidly produce complex components and adds additional development, labor, and material costs. Integrating these steps into the additive manufacturing process chain is critical to maximizing the potential of additive manufacturing [3].Recent changes to the International Organization for Standardization (ISO) Geometric Product Specification (GPS) standards and the American Society of Mechanical Engineers’ (ASME)’s product definition standards have sought to adapt product definitions to reflect technological developments in measurement systems, better relating product definitions to current inspection practices [4], [5]. There are two relatively new ISO GPS standards on linear sizes (ISO 14405-1:2016) and angular sizes (ISO 14405-3:2016) that have been introduced to support engineering requirements whose verification can943only be implemented using a coordinate measurement system (CMS) [6], [7].

The geometric qualification of AM components often necessitates the use of a CMS for feature measurement, as the vast design space allowed by AM enables features that are too complicated for manual gaging techniques. Furthermore, CMSs allow for the automation of complex measurement routines and provide excellent repeatability [8].Several works have sought to characterize the effects of coordinate measurement machine (CMM) measurement parameters on AM components, including investigating the effect of stylus diameter and point sampling strategy [9]–[12]. These were both found to have significant impact on the measurement of AM components. Another influencing factor on measurements using any CMS is feature association. Association is defined in the ISO GPS system as a “feature operation used to fit ideal feature(s) to non-ideal feature(s) according to a criterion” [13].

Association is defined in the ISO GPS system as a “feature operation used to fit ideal feature(s) to non-ideal feature(s) according to a criterion” [13]. An example of this would be fitting a plane to extracted CMM data via a least-squares method to determine the orientation of a surface. While this step in the measurement process is well known for the users of CMSs, it is often overlooked/unknown by those designing AM components. Without the specification of which association criterion to use, the measurement result could vary significantly. For instance, a designer may dimension the diameter of a component intending for it to be measured on a CMM. However, without assigning the association criterion, an inspector would likely only measure the component using calipers or a micrometer, depending on the tolerance requirement. This unnecessary ambiguity can lead to costly mistakes in the production cycle or lead to falsely approved components being used in service.

This unnecessary ambiguity can lead to costly mistakes in the production cycle or lead to falsely approved components being used in service. This work seeks to demonstrate the use of several specification modifiers to convey the required association for a measurement.In this work, AM components are evaluated using manual gaging and a CMM to investigate the effect of the association criterion on the measurement result. The manufacture of the AM components is first presented, including images of the as-built components. The measurement methodologies, including the different association criteria used, are detailed. The details of the individual measurement processes are first presented, followed by a statistical comparison of the different measurement methodologies.

The details of the individual measurement processes are first presented, followed by a statistical comparison of the different measurement methodologies. The results of this comparison indicate no statistical overlap between any of the association criterion chosen, indicating the importance of establishing consistency in this specification to minimize ambiguity in the qualification of AM components.MethodologyThe component investigated in this study was a single volume resembling the reference object described in Ref [14]. The component contains an array of nine 0.5 mm diameter, 2 mm long struts bookended by two 4 mm diameter by 2 mm tall endcaps. Figure 1 shows the component geometry with overall dimensions.

The component contains an array of nine 0.5 mm diameter, 2 mm long struts bookended by two 4 mm diameter by 2 mm tall endcaps. Figure 1 shows the component geometry with overall dimensions. The diameter of the top end cap is investigated in these experiments.944Figure 1: Component investigatedThirty components in two different build orientations with this geometry were produced using an EOS M280 powder bed fusion system with EOS 316L powder at the manufacturer designated parameters for a 20 µm layer height. The components were manufactured in two different orientations to improve the likelihood of success in the build, due to small feature sizes. In the preferred vertical orientation, the component was oriented such that the cylinder axis of the end caps was parallel to the build direction. This orientation was preferred as there should be minimal form error in the struts.

This orientation was preferred as there should be minimal form error in the struts. However, this orientation did have potential for failure, as the small diameter struts were susceptible to deformation during interaction with the recoater blade. Thus, the components were also printed in a horizontal orientation. This orientation had a much higher chance of a successful build, but had a higher potential for form error. Individual barriers were also built around each component to minimize the impact of the recoater blade on the delicate geometry. Figure 2 shows the completed build plate and details of the two different sets of components. One can see in Figure 2(b) that the vertical components did not build successfully. The build failed due to a recoater crash with one of the vertical components. The majority of vertical components appear to have failed prior to this crash.

The build failed due to a recoater crash with one of the vertical components. The majority of vertical components appear to have failed prior to this crash. Since the struts do appear to have been fused, but were plastically deformed, the failure most likely occurred during the first layers of the top end cap. The small layer size deposited on top of loose powder most likely contacted the recoater blade and deformed upward, with the recoater blade dragging it in the direction of travel. For one component, the resistance to deformation must have been high enough to overcome the force of the recoater blade and jammed it in place. Fortunately, all horizontal components completed prior to the build halting, and can be seen in Figure 2(c).Figure 2: Completed AM components before separation from build platform: (a) Build platform layout with vertical oriented components boxed in red and horizontal components boxed in green and unused components in-between.

The direction of recoater travel is shown in the orange arrow. (b) Failed vertical orientation components (c) Successful horizontal orientation componentsThe components were then separated from the build plate using wire electrical discharge machining (EDM). Figure 3 shows an example completed component. The component was imaged using a Leica DVM6 10-megapixel digital microscope in two different orientations: (a) a top down view945of the component (along the build direction), (b) a side view (orthogonal to the build direction). An initial X-ray computed tomography scan of the component was completed and views identical to the digital microscopy were captured. These can be seen in Figures 3(c) and 3(d). There is a stark contrast in the apparent quality of the print between the top down and the side views. In the top view, the component appears to have been manufactured successfully judging by the straight cylindrical struts.

In the top view, the component appears to have been manufactured successfully judging by the straight cylindrical struts. However, significant form error can be seen in the downward facing surfaces in Figures 3(b) and 3(d).Figure 3: Optical and XCT images of a completed AM component: (a) and (c) show a top down view of the component using an digital microscope and XCT reconstruction respectively, (b) and (d) show the side view of the component with the build direction indicatedUnfortunately, three components were damaged during the wire EDM process, so 27 components were used in the remaining measurement processes. The components were first measured individually using a calibrated micrometer, which had a measurement resolution of 0.0025 mm. The components were then fixtured and measured individually on a CMM.

The components were first measured individually using a calibrated micrometer, which had a measurement resolution of 0.0025 mm. The components were then fixtured and measured individually on a CMM. The CMM measurements were completed using a Zeiss Micura with a maximum permissible error of length measurement (E0,MPE) of (0.8 + L/400) μm, where L is the measured length in mm, according to ISO 10360-2 [15]. The CMM measurements used a 50 mN probing force, a 0.5 mm probe tip diameter, and thermal compensation active using a coefficient of thermal expansion of 16.2 μm/(m·K). The outer diameter of the top endcap was measured by probing 475 points on the surfaces.

The outer diameter of the top endcap was measured by probing 475 points on the surfaces. The fixtured component and the sampled points, shown as the green arrows, in the measurement can be seen in Figure 4(a) and (b), respectively.Figure 4: (a) Component CMM measurement setup and (b) points sample in CMM measurementAfter the data was captured using the CMM, it was imported into GOM Inspect software [16] as a point cloud and the diameter of the feature was calculated using different association criteria. These different association methodologies are depicted in Figure 5, using the nomenclature defined in ISO94614405-1:2016 [6]. Commonly, dimensions of components in product definition data are often listed without modifiers, as in Figure 5(a). This then implies the default specification of two-point size. A two- point size is the distance separating the two points composing an opposite point pair taken simultaneously on the extracted feature (see ISO 17450-3) [17].

A two- point size is the distance separating the two points composing an opposite point pair taken simultaneously on the extracted feature (see ISO 17450-3) [17]. A two-point measurement also can be specifically designated using the (LP) modifier, as in Figure 5(b). Two-point measurements are representative of mechanical gauging processes, such as measurements with calipers, micrometers, and comparators, where a measurement is taken across two opposing parallel faces. However, measurements executed using a CMM often do not use this measurement specification. CMM measurements are often executed by extracting data points from the surface and fitting features using a least-squares criterion. The least-squares association (or global gaussian) can be designated by using the (GG) modifier, as shown in Figure 5(c). Another common association criterion used is the Chebyshev (or minimax), which can be designated using the modifier (GC), as shown in Figure 5(d).

Another common association criterion used is the Chebyshev (or minimax), which can be designated using the modifier (GC), as shown in Figure 5(d). An additional commonly used association criterion for cylindrical bosses is the minimum circumscribed cylinder, which can be specified using the modifier (GN), as in Figure 5(e).Figure 5: Drawing depiction of measurement specification modifiers used for the different association criterion: (a) Default (two-point measurement), (b) Two-point measurement, (c) Least-squares, (d) Chebyshev, (e) Minimum circumscribed cylinderThe AM components were assessed using all of the specification modifiers shown in Figure 5. The no modifier case and (LP) specification are identical and were assessed using the micrometer measurements. The (GG), (GC), and (GN) specifications were assessed by changing the fitting algorithm used on the extracted measurement data for each sample. This then created four unique diameter measurement results for each sample.

This then created four unique diameter measurement results for each sample. Statistical analysis was then performed on the measurement results to discern differences in the measured diameter. Initial normality tests were conducted on each of the measurement specifications using an Anderson-Darling test using the null hypothesis of normality and a confidence value of α = 0.05. An analysis of variance (ANOVA) was then conducted with the null hypothesis that there is no significant difference between the means of the measurement methods with a confidence value of α = 0.05. If the null hypothesis was rejected, a Tukey- Kramer pairwise comparison was then conducted to determine if there were any overlaps in the confidence intervals of the various groups with a confidence value of α = 0.05 to test if the means of the two data sets compared are very similar.ResultsFigure 6 shows the results of the micrometer diameter measurements of the components.

As previously described, these measurements used a two-point measurement technique to measure the diameter. The AM components appear, on average, within the specification. The average diameter for947these components was measured to be 4.014 mm, with a standard deviation of 14 µm. The left side of Figure 6 shows the histogram of the measured values. The vertical black dashed line indicates the mean of the measured values. The vertical red dashed line indicates three standard deviations away from the mean, indicating close to where the limits of a normal distribution would lie. Examining this, the data does appear to follow a normal distribution. The right side of Figure 6 displays the quantile- quantile (q-q) plot using a Gaussian distribution to analyze the normality of the diameter measurements.

The right side of Figure 6 displays the quantile- quantile (q-q) plot using a Gaussian distribution to analyze the normality of the diameter measurements. The Anderson-Darling test failed to reject the null hypothesis of normality and the data appears to closely follow the normal distribution with the exception of two outliers at approximately 3.99 mm.Figure 6: Diameter measurement using the micrometer, (LP) two-point size modifierFigure 7 displays the results of the CMM measurements using the least-squares association criterion (GG). The average for these measurements is 3.989 mm. This is in contrast to the two-point measurements conducted with the micrometer, as the two means are 25 µm apart. The standard deviation of these measurements was calculated to be 16 µm. However, unlike the two-point measurements, the histogram of the measurements does not appear to follow a normal distribution, as the measurements are skewed toward larger diameters.

However, unlike the two-point measurements, the histogram of the measurements does not appear to follow a normal distribution, as the measurements are skewed toward larger diameters. The Anderson-Darling test rejects the null hypothesis that the measurements come from an underlying normal distribution. Significant departures from the underlying distribution can be seen in the q-q plot. This confirms that two different measurement techniques performed on AM components can not only shift the mean of the measured values, but also yield significantly different underlying distributions of the results.Figure 7: Diameter measurement using the leas-squares association criterion on the CMM data, (GG) modifier948Figure 8 displays the results of the CMM measurements using the Chebyshev association criterion (GC). The average for these measurements was 4.029 mm, which is larger than both the LP and GG measurement methodologies.

The average for these measurements was 4.029 mm, which is larger than both the LP and GG measurement methodologies. The standard deviation of these measurements was 25 µm, which is also larger than the two previously discuss methodologies. The distribution of the data is clearly one sided toward larger diameter measurements. As expected, the null hypothesis of normality is rejected in the Anderson-Darling test and the q-q plots show poor conformance to the underlying distribution throughout, not only at the tails of the distribution.Figure 8: Diameter measurement using the Chebyshev association criterion on the CMM data, (GC) modifierFigure 9 displays the results of the CMM measurements using the minimum circumscribed cylinder association criterion (GN). The average of these measurements is significantly larger than the other measurement methods, 4.139 mm. The dispersion of data is also larger, with a standard deviation of 35 µm.

The average of these measurements is significantly larger than the other measurement methods, 4.139 mm. The dispersion of data is also larger, with a standard deviation of 35 µm. The histogram of the measurement results appears gaussian. This is confirmed by the Anderson-Darling test which fails to reject the null hypothesis that the data comes from an underlying normal distribution. The q-q plot shows some minor departures from the underlying distribution around the tails, but the majority of the data appears to conform well. The (GN) data by far indicates the largest diameters measured. This is not unexpected as the minimum circumscribed feature is only influenced by the highpoints on the surface. However, it is important to note that the minimum circumscribed criterion generally does not give an ideal normal curve from the theoretical perspective, but provides a truncated normal distribution bounded by the maximum deviation of the measurement or manufacturing process [18].

A non-truncated normal distribution is used here for simplicity and for comparison.949Figure 9: Diameter measurement using the minimum circumscribed association criterion on the CMM data, (GN) modifierTable 1 displays the results of the ANOVA between the four different measurement procedures. Clearly, the null hypothesis that the four data sets all come from one underlying distribution is rejected. Figure 10 displays a box plot comparing the four different data sets. The red line in each data set represents the median value. The top and bottom edges of the blue box indicate the 25th and 75th percentiles of the data, respectively. The whiskers extend to the most extreme data points not considered outliers, and the outliers are plotted individually using the '+' marker symbol. The (GN) data is clearly larger than the other three data sets showing no overlap at all with the exception of one outlier in the (GC) data set.

The (GN) data is clearly larger than the other three data sets showing no overlap at all with the exception of one outlier in the (GC) data set. This separation clearly contributed to the rejection of the null hypothesis in the ANOVA. The other three measurement methods are much closer and have overlap at the extreme values. However, the (GG) measurements do appear lower than the rest, but it is difficult to discern from these results.

However, the (GG) measurements do appear lower than the rest, but it is difficult to discern from these results. It is important to note that one of the underlying assumptions of an ANOVA is the normality of data, but it has been shown that, generally, ANOVA is robust to violations of the normality assumption [19].Table 1: ANOVA results for the comparison of the association criterionSourceSSdfMSFP-ValueGroups0.355830.1186388.228.92e-47Parts0.0355260.00144.471.45e-07Error0.0238780.0003Total0.4152107Figure 10: Box plot comparing the different measurement methods950Table 2 shows the results from the Tukey-Kramer pairwise comparison test. Each difference column indicates the difference between the mean of the “Group” and “Control Group” in the individual comparison. The P-Values indicate the results of the comparison and show no overlap between the six pairs tested at a confidence value of α = 0.05.

The P-Values indicate the results of the comparison and show no overlap between the six pairs tested at a confidence value of α = 0.05. Figure 11 graphically displays the results of the test. The circles for each data set indicate the mean value, while the horizontal lines indicate the 95% confidence interval for each data set. It is clear that none of the six pairs investigated have statistical overlap, as reflected in the P-Values of the test. Similar to ANOVA, the Tukey-Kramer analysis has been shown to be relatively robust to non-normality, but Type I and Type II errors are certainly possible [20]–[22].

Similar to ANOVA, the Tukey-Kramer analysis has been shown to be relatively robust to non-normality, but Type I and Type II errors are certainly possible [20]–[22]. Clearly, each of these different association criteria has had a significant impact on the diameter measurement of these AM components.Table 2: Results from Tukey-Kramer pairwise comparisonGroupLPControl Group GGLower Limit (mm)0.0073Difference (mm) 0.0243Upper Limit (mm) 0.0413P-Value0.0LPGC0.03250.01550.00150.0088LPGN0.14250.12550.10850.0GGGC0.05680.03980.02280.0GGGN0.16680.14980.13280.0GCGN0.12700.11000.09300.0Figure 11: Graphical display of the Tukey-Kramer analysisIt is important to note that the magnitudes and statistical findings of the results in this work are specific to the components analyzed.

While measurements using the (GG) method will likely be smaller than measurements using the (LP) and (GN), the deviations between the results may not always be at the magnitude observed in this work. The values will depend on the component size, component form, and surface texture. An increase in form errors will directly impact the (GN) measurements and will likely impact the (LP) measurements, since these methods are heavily influenced by the extreme values of the surface. However, these errors will have a lesser impact on the (GG) method since the least- squares method seeks to minimize the error from all points to the associated feature. For instance, if the axis of the cylinder had been aligned along the build axis, as described in the methodology, we may have seen very different results as the form errors on the cylinder surface would be less likely to occur.

Furthermore, if the components were larger, component form would also change, as a greater number951of layers would be able to form the cylindrical geometry. Future work should investigate the effect of varying component geometry and build orientation on these measurement methodologies.These results show that significant differences in the measurement of AM components can occur depending on the measurement methodology implemented, specifically the choice of association criterion. It is critical for designers, manufacturers, and inspectors of AM components to understand the differences in these methodologies and their potential impact on the measurement result. Designers should seek to provide specifications using the intended inspection methodology, specifically by using the specification modifier appropriate for their application, in order to minimize ambiguity in the product requirements.

Likewise, this onus also falls on the inspector to report the measurement procedure used to produce the results.ConclusionsThis work examined the use of different measurement methods in the qualification of diameter on a simple AM feature. The results from this investigation showed that the choice of the association criterion can have a significant impact on the measurement result. The use of different association criteria was found to produce different mean values, different dispersions of values, and different distributions of data across the measurement of 27 components. The different association criteria were found to produce statistically significant measurement results, with differences in means up to about 0.15 mm.

The different association criteria were found to produce statistically significant measurement results, with differences in means up to about 0.15 mm. Further investigations should examine the impact of these specifications over a wider range of geometric features, sizes of features, and orientations within the AM build volume to gain a better understanding of the potential impact of association criterion on the measurement of AM components.DisclaimerCertain commercial entities, equipment, or materials may be identified in this document to describe an experimental procedure or concept adequately. Such identification is not intended to imply recommendation or endorsement by the National Institute of Standards and Technology, nor is it intended to imply that the entities, materials, or equipment are necessarily the best available for the purpose.References[1] “Roadmap for Automotive Additive Manufacturing,” United States Council for Automotive Research LLC, 2021. [2] R. K.

viewed PaperIMPACT OF MOISTURE ABSORPTION ON 3D PRINTING NYLON FILAMENTHaijun Gong, Michael Runzi, Zezheng Wang, Lianjun WuDepartment of Manufacturing Engineering, Georgia Southern University, Statesboro, GA 30460AbstractNylon is a commonly used thermoplastic material featured for its high resistance to heat, chemical, and impact. Nylon filaments that facilitate complex part manufacturing through the extrusion-based 3D printing process are widely used in industry and can be easily acquired from various vendors or suppliers. However, as a hygroscopic material, nylon directly absorbs moisture from the surrounding air at a fast rate. This study concentrates on evaluating the moisture absorption of nylon filament in a humid environment. Experiments are conducted to correlate the moisture content of nylon filament with its dwelling time in a humidity chamber. Test artifacts are printed to demonstrate the influence of moisture on nylon part quality.

Test artifacts are printed to demonstrate the influence of moisture on nylon part quality. Mechanical testing is also carried out for analyzing the material property degradation due to increased moisture content. Suggestions on nylon filament usage and storage are discussed.IntroductionAs a commonly used thermoplastic material, nylon designates a family of synthetic polymers composed of polyamides of high molecular weight. Nylon polymers are tough, strong, and outstanding for its high resistance to heat, chemical, and impact, which are of great interest for commercial applications. With the introduction of additive manufacturing or 3D printing to the market, nylon becomes a popular feedstock material in the form of powder (for selective laser sintering process) and filament (for fused deposition modeling process). Especially, 3D printing nylon filament affords great convenience for making parts of low-volume production or complex structures, using an extrusion-based 3D printer.

Especially, 3D printing nylon filament affords great convenience for making parts of low-volume production or complex structures, using an extrusion-based 3D printer. Also, nylon filaments are low-cost and easily acquired through various vendors or suppliers, making 3D printing for nylon parts very accessible to end users.The influence of moisture on 3D printing is of great interest in a wide range of areas. Many commonly used 3D printing filaments, such as polylactic acid (PLA), acrylonitrile butadiene styrene (ABS), polycarbonate, or related composites, were studied revealing how the material properties are changed and how the impact of moisture is addressed [1-4]. Some 3D printing filaments for pharmaceutical application were investigated for stability challenges of varied storage conditions including temperature and relative humidity [5]. Shrinkage behavior of building materials for 3D printing was also studied regarding the different drying or moistening conditions [6].

Shrinkage behavior of building materials for 3D printing was also studied regarding the different drying or moistening conditions [6]. As a hygroscopic material, nylon plastic can absorb moisture from surrounding humidity. Monson et al [7] reported the dimensional changes associated with moisture absorption of various polyamides. Banjo et al [8] 3D printed nylon specimens and immersed in deionized water at different temperature and concluded that nylon-based materials absorb much more water than 3D printing PLA specimens. A direct relationship between moisture absorption and reduction in flexural properties was also observed for the nylon specimens. The degradation of nylon material16can be attributed to the amount of filament moisture, which acts as a plasticizer reducing the molecular entanglements and bonding between the molecules [9].

The degradation of nylon material16can be attributed to the amount of filament moisture, which acts as a plasticizer reducing the molecular entanglements and bonding between the molecules [9]. It has been found that water sorption (immersion into liquid water bath) has very different transport mechanism at surface compared to the moisture sorption (moisture in air) [10]. Thus, this project extends to study the effects of humidity on the feedstock filament of nylon materials in the 3D printing process. The characteristics of 3D printed nylon specimens are evaluated at various moistening time (moisture contents) to reveal the influence.ExperimentThe feedstock material OVERTURE Easy Nylon filament (1.75 mm) was acquired for this study. The virgin filament was placed in a MEMMERT humidity chamber for absorbing moisture. The humidity chamber was set with a constant temperature (40 °C) and humidity (80 % rh) to mimic a stable humid environment.

The humidity chamber was set with a constant temperature (40 °C) and humidity (80 % rh) to mimic a stable humid environment. After a certain amount of time, the moisture content of the nylon filament was measured using a Torbal moisture analyzer (model no. BTS110). The nylon filament was then fed into an extrusion-based CreatBot F430 3D printer for printing specimens. Three types of nylon artifacts were created for comparison of varied moisture content: extruded material, thin-wall samples, and tensile testing specimens. The extruded material is the continuous nylon fiber being pushed out of the nozzle. The thin-wall sample has a hollow cubical shape with a wall of stacked nylon strand. The size of the tensile testing specimens is in compliance with the ASTM D638 standard (tensile testing for plastics), with a thickness 4 mm. The tensile specimens are printed with a YZ orientation, as per ISO/ASTM 52921 standard.

The tensile specimens are printed with a YZ orientation, as per ISO/ASTM 52921 standard. The extruded materials and thin-wall samples were observed using an Olympus optical microscope (BX53M) and a Keyence 3D Measurement Macroscope (VR-3200). Tensile specimens were tested using a MTS CriterionTM Model 43 tensile tester at room temperature. The experimental procedure of moistening and 3D printing is illustrated in Fig. 1.Fig. 1 Experimental set upResults and DiscussionMoisture measurementAs a hygroscopic material, nylon directly absorbs moisture from the surrounding air at a fast rate. In this study, nylon filament was placed in the humidity chamber by varying the dwelling time at four levels (0h, 24h, 48h, and 72h). Time=0h represents virgin filament without any moistening. The measured moisture content of nylon filament is shown below in Fig. 2.176)%5. tw(t n e t n o c43e r u t s i o M2100244872Dwelling Time (h)Fig.

The measured moisture content of nylon filament is shown below in Fig. 2.176)%5. tw(t n e t n o c43e r u t s i o M2100244872Dwelling Time (h)Fig. 2 Moisture versus dwelling time of nylon filament in the humidity chamberThe plotted curve in Fig. 2 exhibits a dramatic increase in moisture content with the extended dwelling time. It is noted that even the virgin nylon filament (0h) contains about 1 wt.% water, which is ascribed to the moisture absorption during the filament production. The hygroscopicity of nylon material causes an inevitable moisture absorption when a package of nylon filament is installed for an extrusion-based 3D printer, especially for a printer without full enclosure. When the environmental humidity reaches to a high level, the nylon filament continuously absorbs the moisture in the air. This can be demonstrated by the moisture curve illustrated in Fig. 2.

When the environmental humidity reaches to a high level, the nylon filament continuously absorbs the moisture in the air. This can be demonstrated by the moisture curve illustrated in Fig. 2. It can be seen that, in a warm and humid environment (40 °C and 80% rh), the moisture content of nylon filament is proportional to the dwelling time. The moisture of nylon filament is measured to be around 5.5 wt.% after being placed in the humidity chamber for three days (72h), showing no sign of saturation. The nylon filament exhibits comparable absorption rate to the nylon samples immersed in deionized water (Ref [8]).Extruded MaterialExtrusion-based 3D printing process, such as FDM or FFF, extrudes melted material (like a fiber) in thin strands and bonds to the prior layers or substate. In this study, the extruded material is sampled by collecting the nylon fibers of different moistening time. As shown in Fig.

In this study, the extruded material is sampled by collecting the nylon fibers of different moistening time. As shown in Fig. 3, the morphology and consistency of the extruded material are significantly influenced by the moisture content of the feedstock filament. Generally, 0h and 24h nylon fibers maintain a consistent shape of circular cross-section about 0.4mm in diameter. A few bubbles originated from water evaporation burst after nylon material is extruded out of the nozzle tip. Hence, the defective sites are observable on the surface of the 24h nylon fiber. However, if a high moisture content is introduced in the nylon filaments (48h and 72h), the evaporated water bubbles severely deteriorate the nylon fiber morphology, due to the interior bubble expanding and exterior bubble bursting. The viscosity of melted nylon decreases with the increased moisture uptake, which expedites the void formation or foaming.

The viscosity of melted nylon decreases with the increased moisture uptake, which expedites the void formation or foaming. Also, the formation of hydrogen bonds between water and carbonyl and ether linkages in the polymer chain causes the “pseudo-crosslinking” effect [11], which results in the extruded material degradation. A smooth and straight nylon fiber cannot be achieved. Some18expanded big bubble even disrupts the continuous fiber and causes an extrusion inconsistency. The high moisture content is believed to negatively influence the 3D printing parts of nylon filament.0h24h48h72h1 mmFig. 3 Schematic of extrusion-based 3D printing process and the extruded nylon materials of different moistening timeSurface finishFour thin-wall samples were printed using the moistened filaments to reveal the influence of moisture on surface finish of 3D printing nylon part. The lateral wall is characterized using the Keyence automatic 3D measuring macroscope.

The lateral wall is characterized using the Keyence automatic 3D measuring macroscope. It is found that the surface quality is dramatically decreased, ascribing to the deteriorated nylon strand with a higher moisture. Fig. 4 clearly visualizes the change of surface finish with specific moistening time (or increased moisture content) of nylon filaments.0h24hThin-wall sample48h72h3 mmFig. 4 Lateral surface profile of thin-wall nylon samplesThe surface roughness Ra was evaluated based on the scanned surface profile. The evaluation length is ~5mm and directed along the building direction (orthogonal to the layers).19Periodic peak and valley can be observed on the roughness profile, as shown in Fig. 5. The Ra value is automatically computed through the profile heights over the evaluation length. It is noted that the Ra value does not change as much as observed surface finish, even though the roughness profile of 48h and 72h samples presents apparent irregularity.

It is noted that the Ra value does not change as much as observed surface finish, even though the roughness profile of 48h and 72h samples presents apparent irregularity. The influence of moisture on the surface quality of 3D printing nylon parts can be easily disregarded if only Ra measurement is conducted. Visual comparison to the nylon surface finish of various moisture contents is usually not achievable if the nylon filament has already been moistened. 3D printing part evaluation is suggested for multiple metrological methodologies to confirm its surface quality.(a) 0hRa 15μm(b) 24hRa 22μm(c) 48hRa 16μm(d) 72hRa 16μmFig. 5 Roughness profile of nylon thin-wall samplesMechanical testingTensile test was conducted at room temperature to specimens which were 3D printed using the virgin nylon filament (0h) and moistened nylon filament (72h). The test aimed to demonstrate the effect of moisture content on the tensile strength of nylon materials.

The test aimed to demonstrate the effect of moisture content on the tensile strength of nylon materials. Thus, only two types of filaments, with very distinct moisture content (0.9% vs. 5.5%), were adopted. The tensile specimens conform to the ASTM D638 Type I size requirement, with a thickness of 4 mm. Four duplicates were tested for each type of tensile specimen. The stress-strain curve is shown in Fig. 6. 3D printing plastic materials usually present a varied mechanical property. Hence, it is anticipated the slight variation of the tensile curves occurs between 0h nylon specimens, such as elastic modulus (1.50~1.81 GPa). It is noted that two 0h tensile specimens were not ruptured during the test because the extensometer reached its strain limitation. Two stress-strain curves of 72h specimens are completely overlapped. Overall, the nylon specimens which were 3D printed using virgin filament exhibit great elasticity and ductility.

Two stress-strain curves of 72h specimens are completely overlapped. Overall, the nylon specimens which were 3D printed using virgin filament exhibit great elasticity and ductility. However, the tensile properties of the nylon specimens printed using 72h moistened filament are clearly degraded, especially in ductility. The max strain value of 72h tensile specimens is only 8%, which is significantly lower than the 0h20specimens. The yield strength and ultimate tensile strength are also lowered due to a high percentage of moisture content. The degraded nylon material of extrusion is believed to contribute to the inferior tensile property. Aforementioned, the void formation and “pseudo-crosslinking” effect cause the extruded nylon material degradation. Moisture acts as a plasticizer and reduces the strength and stiffness of the nylon material.

Moisture acts as a plasticizer and reduces the strength and stiffness of the nylon material. In addition, water molecules diffuse into the nylon structure and bond to the polymer chain within the amorphous regions, forcing the polymer chains apart. This would substantially weaken the layer bonding during 3D printing nylon filaments. Overall, these factors are composited to cause the decreased ductility of nylon specimens printed with filament being exposed in the moisture.6050Not ruptured40) a P M30(s s e r t S20Tensile specimens printed by 0h filament 72h filament10000.050.10.150.2Strain (mm/mm)Fig. 6 Stress-strain curve of tensile testing specimens of nylonThe influence of moisture on 3D printing nylon filament, not limited to nylon material, deserves extra attentions, based on the results and discussions of this study.

It is noted that the quality of material extrusion, surface finish, and mechanical property are all diminished due to the moisture absorbed by the nylon filament. As a hygroscopic material, the usage and storage of nylon filament must be carefully handled and managed to ensure to the 3D printing part quality. The filament operation guidelines and protocols are suggested to be proposed to avoid the inconsistency of mechanical properties of 3D printing parts. In general, the environmental humidity must be well controlled when using nylon filament for 3D printing. An enclosed chamber is recommended for the printing process. The unused nylon filament should be wrapped in a vacuum condition to avoid contacting moisture in the air. If the filament is out of supervisory monitoring for a long term, the moisture analysis should be conducted to verify its eligibility for quality part.

If the filament is out of supervisory monitoring for a long term, the moisture analysis should be conducted to verify its eligibility for quality part. Otherwise, the suitable drying must be performed before reusing the moistened filament.210.25ConclusionThis study relates the moisture content of 3D printing nylon filament with the dwelling time in a constant humid environment. The moisture measurement confirms the hygroscopicity of nylon filament. The negative influence of moisture on nylon filament is demonstrated through analyzing extruded materials, surface finish, and tensile property of nylon specimens. It is found that the extruded nylon material is significantly degraded if the filament is highly moistened, causing an irregular surface finish as characterized. The tensile property is also lowered, especially for ductility, due to the composited factors of water molecules to the nylon filament.

viewed PaperImpact of Nozzle Condition on Powder Catchment Efficiency for Coaxial Powder Direct Energy DepositionLauren Heinrich1,2*, Joseph Fletcher1, Thomas Feldhausen2, Thomas Kurfess1, Christopher Saldaña11 Georgia Institute of Technology, George W. Woodruff School of Mechanical Engineering, Atlanta, GA 30332, USA 2Oak Ridge National Laboratory, Manufacturing Science Division, 1 Bethel Valley Rd, Oak Ridge, TN 37830, USACorresponding author: lheinrich6@gatech.eduAbstractBlown powder directed energy deposition is well-designed for fine resolution additive manufacturing processing. Coaxial powder deposition heads use an outer layer of shielding gas directed by an outer nozzle to prevent oxidation occurring during the powder melting process. Powder blown feedstock catchment efficiency can be as low as 50-80% whereas wire deposition systems are closer to 98% efficient.

Powder blown feedstock catchment efficiency can be as low as 50-80% whereas wire deposition systems are closer to 98% efficient. The present study evaluates the impact of directed energy deposition nozzle condition on catchment efficiency. Changes in the overall outer shielding gas nozzle length has been found to increase material usage efficiency by 10% through convergence of the powder flow. The results of this experiment show that for coaxial powder deposition head design, if the standoff distance can safely be decreased, powder catchment efficiency can be increased as the outer shielding gas nozzle is increased in length, or the standoff distance is decreased.Keywords: Hybrid manufacturing, Blown-powder DED, Catchment efficiencyIntroductionAdditive manufacturing continues to be a topic of research as the freeform flexibility with manufactured geometry for low volume needs such as aerospace, medical, and industrial applications have shown to be useful [1].

Direct energy deposition (DED) is a form of additive manufacturing such as wire arc additive manufacturing, wire fed laser melt deposition, or blown powder laser melt additive manufacturing [2]. DED is profitable when high buy-to-fly components are manufactured due to the ability to add material in the net shape of the final component; however, the surface finish is often not adequate for many uses [3]. Hybrid machine tools were developed to meet the need for selective material addition through DED combined with the precision machining of the native machine tool in one machine platform [4–6]. This investigation evaluates the efficiency of blown powder DED additive manufacturing processing on a hybrid system where the powder was deposited using a coaxial deposition head with a center laser rated at a power output of up to 1 kW.1091Powder DED often has low catchment efficiency ranging from below 50% up to 80% whereas wire systems are capable of efficiencies greater than 98% [7].

Catchment efficiency is the amount of material adhered to the deposition versus the total amount of feedstock used. An investigation was performed to determine if powder from a DED system could be reused, and Terrassa et al. found the material should be sieved each time and the amount of oxygen in the feedstock increased after each re-use, but recycling powder was effective [8]. This experiment was performed in a machine tool that did not utilize coolant which contaminates the powder to a higher degree. Coolant is critical to assist with the lubrication, corrosion inhibition, chip evacuation, and heat removal during subtractive machining processing [9]. Many hybrid machines will have the powder contaminated with coolant, so a different strategy is needed to increase powder usage efficiency.

Many hybrid machines will have the powder contaminated with coolant, so a different strategy is needed to increase powder usage efficiency. Currently, powder is often disposed of according to local disposal laws [10].Many studies have been performed to evaluate the effectiveness of various strategies to improve material catchment efficiency. Haley et al. found that there was an ideal deposition standoff distance between the powder nozzles and the deposition location which was located within the powder convergence point [11]. It was also found that there was a passive feedback loop with the deposition head employed which stabilized the working distance while sacrificing powder efficiency [11]. It was found that by increasing the angle of the powder flow and decreasing the convergence distance increases catchment efficiency [12]. Smith et al.

It was found that by increasing the angle of the powder flow and decreasing the convergence distance increases catchment efficiency [12]. Smith et al. evaluated the effectiveness of magnetizing the build plate to increase powder catchment near the melt pool but found that there were often inconsistencies in the build height due to the inconsistency in the magnetic flux [13]. Powder spatter is a reason for large powder loss, partly due to powder rebounding off an unmelted powder island and off the molten melt pool [14,15]. Even with these experiments of powder catchment efficiency, traverse feed rate, powder mass flow rate, laser power, gas flow, and pre-heat must be optimized for increased powder catchment and deposition quality [16–19].

When these parameters are optimized for the given DED system, the values can be scaled to increase the deposition rate while retaining a similar energy density if adequate thermal energy dissipation is possible [20,21].This study aims to investigate the effectiveness of a more universal approach to increasing powder catchment efficiency for any blown powder DED system by only changing either the standoff distance or outer shielding gas nozzle length (hereon referred as the nozzle). This mitigates the need to redesign the entire deposition head. It was found that powder efficiency decreased from 55.5% to 46% when the nozzle length was decreased by 2.47 mm and the mass flow and standoff distance was kept consistent as was found in experiment 1. The catchment efficiency was increased by nearly 6% by increasing the mass flow and achieving consistent bead height throughout the deposition process as was found in experiment 2.

The catchment efficiency was increased by nearly 6% by increasing the mass flow and achieving consistent bead height throughout the deposition process as was found in experiment 2. The catchment efficiency was increased by nearly 10% by decreasing the standoff distance to the powder convergence point as was found in experiment 3.1092Experimental ProcedureAll deposition samples were fabricated using a Mazak VC500-AM 5-axis CNC machine equipped with a dual hopper powder feeder. Argon was used as the carrier and shielding gas for the deposition head. The deposition head featured a 1-kW IPG Photonics YLS 1000 fiber optic laser with a 1.2 mm spot to melt the feedstock. 410 stainless steel powder from JAS was utilized as the feedstock with a powder size of +45-150 µm. The CAM software, hyperMILL by Open Mind was used to design the test specimen and generate the required G-code.

The CAM software, hyperMILL by Open Mind was used to design the test specimen and generate the required G-code. The single bead width deposition geometry for all experiments was a circular tube with bead centerline of 30 mm and 15.6 mm tall as can be seen in Figure 1. A 5-axis continuous helical toolpath with a 0.59 mm pitch and feed rate of 250 mm/min was used to deposit samples onto a 304 stainless steel build plate. The deposition toolpath can be seen in Figure 1.Figure 1. CAD/CAM 5-axis helical tool path at 0.59 mm pitch utilized for each sample where (a) isometric and (b) deposition simulationFour different nozzles with lengths relative to the original equipment manufacturer’s length of +1.47 mm. +.59 mm, 0 mm, and -1 mm were interchanged to analyze the effects of standoff distances, mass flow rates, and catchment efficiency. The original equipment manufacturer’s nozzle is flush with the powder nozzle as can be seen in Figure 2.

The original equipment manufacturer’s nozzle is flush with the powder nozzle as can be seen in Figure 2. Laser power of 430 W was kept constant throughout all trials which was determined for consistent build quality from testing with the longest nozzle. The deposition substrate was 304 stainless steel that was 6.25 mm thick and 76 mm in length and width where the top and bottom surfaces were machined to remove the mill scale. Each build plate was weighed before and after deposition to determine the added mass from the deposition process. A 2-minute powder mass flow test was performed before and after each deposition to ensure consistent mass flow results of the powder feeding system. The laser fire time during the deposition of the tube geometry was recorded to determine how long the powder was fed during the deposition process.

The laser fire time during the deposition of the tube geometry was recorded to determine how long the powder was fed during the deposition process. The shortest point on the build was measured as to determine the useful height if a machining process was to be used to finish the sides and top to a consistent geometry.1093Figure 2. Blown powder coaxial DED nozzle assembly diagramThe first experiment evaluated the catchment efficiency of various nozzle lengths. The control mass flow rate was determined by the +1.47 mm nozzle which produced a tube within a 3% tolerance of the expected final height of 15.6 mm and a mass flow of 3.5 g/min. The second experiment focused on the catchment efficiency for each of the different nozzles when the mass flow rate was increased to produce a tube within tolerance of the programmed dimensions. The third experiment investigated the effect of standoff distance on powder catchment for various nozzle lengths.

The third experiment investigated the effect of standoff distance on powder catchment for various nozzle lengths. The results of the experiments will be discussed in the following section.Results and DiscussionThe first experiment was used to determine the decrease in catchment efficiency as the shielding gas nozzle length was decreased and all other deposition parameters were kept constant as seen in Table 1. The standoff height was kept at nominal distance of 6mm to ensure equal spot size. The catchment efficiency decreased from 55.5% with the 1.47 mm long nozzle down to 46.0% with the -1.0 mm nozzle as seen in Table 1. These results are consistent with the idea that as the deposition pool deviates further from the convergence point, deposition efficiency decreases [11]. This also illustrates the effect of compounding error when the layer height is programmed higher than the effective deposition bead height.

This also illustrates the effect of compounding error when the layer height is programmed higher than the effective deposition bead height. The stability of the process diverges, causing decreased material efficiency as the process continues at the consistent deposition parameters. Table A1 in the Appendix shows images of the resulting part from each experiment.1094Table 1. Experiment 1: Catchment efficiency results for varied shielding gas nozzle lengthWith decreasing nozzle length, it is hypothesized that this change is affecting the convergence distance of the powder flow. This is strengthened by the observation of decreasing catchment efficiency as nozzle length is decreased. Ongoing work is being conducted using computational fluid dynamics to further understand how the nozzle length affects the powder convergence.

Ongoing work is being conducted using computational fluid dynamics to further understand how the nozzle length affects the powder convergence. Building on this initial experiment, experiment 2 evaluates how increasing the mass flow can compensate for the decreased final build heights observed in experiment 1.The second experiment focused on the catchment efficiency for each of the four different shielding gas nozzles when the mass flow rate was increased to produce a tube within 5% tolerance of the 15.6 mm programmed target height. The results of the experiment can be seen in Table 2. While the mass flow was increased by less than 7% from the longest to the shortest nozzle evaluated, the deposition efficiency was kept within 10% between all depositions. This experiment illustrates that if a consistent bead height over the duration of the build time can be achieved, deposition catchment efficiency can be improved.Table 2.

This experiment illustrates that if a consistent bead height over the duration of the build time can be achieved, deposition catchment efficiency can be improved.Table 2. Experiment 2: Catchment efficiency results for increased mass flowInstead of increasing the mass flow to compensate for decreased build height, another strategy is to decrease the standoff distance. By decreasing the nozzle standoff distance, the issue of reduced powder catchment may be able to be mitigated. If true, this further strengthens the hypothesis of an increased nozzle length decreasing the powder convergence distance. It has been found that by depositing at the convergence point can increase deposition efficiency [11]. To test this, the third experiment compares the catchment efficiencies of the two extreme nozzle lengths1095with different standoff distances.

To test this, the third experiment compares the catchment efficiencies of the two extreme nozzle lengths1095with different standoff distances. Decreasing the standoff distance decreases the spot size as well as the clearance between the nozzle and the deposited material increasing chances of colliding the nozzle into the workpiece. These standoff distances were determined by preliminary simulation results. In the case of decreasing the standoff distance of the OEM nozzle to that nearly equal to the longest nozzle, equivalent catchment efficiency was achieved as can be seen in Table 3. By decreasing the standoff distance, the powder convergence point was brought closer to the deposition location, consistent bead height was achieved, and material efficiency of the OEM nozzle was equal to that of the longest nozzle.Table 3.

Experiment 3: Catchment efficiency results for decreased standoff distanceIn summary, when greater efficiency is desired and nozzle redesign is prohibitive, decreasing the standoff distance to that of the convergence point is a viable option no matter the powder blown DED deposition head type as was found in experiment 3. If the interference of the head is a concern, a modified nozzle with an increased shielding gas nozzle length to help direct the powder more controlled to the deposition location further from the nozzle is the next most efficient strategy as was found in experiment 1.

Lastly, if standoff distance cannot be decreased and increasing the nozzle length is not possible, standoff distance can be increased while also increasing mass flow in order to achieve a steady bead size relative to the programmed bead height with higher material efficiency as was found in experiment 2.Summary and ConclusionsThis study evaluated various ways to increase powder catchment of a blown powder DED deposition head through varied shielding gas nozzle lengths, mass flows, and standoff distances. It was found that powder efficiency decreased from 55.5% to 46% when the nozzle length was decreased by 2.47 mm and the mass flow and standoff distance were kept consistent as was shown in experiment 1. The catchment efficiency was increased by nearly 6% by increasing the mass flow and achieving consistent bead height throughout the deposition process as was shown in experiment 2.

The catchment efficiency was increased by nearly 6% by increasing the mass flow and achieving consistent bead height throughout the deposition process as was shown in experiment 2. The catchment efficiency was increased by nearly 10% by decreasing the standoff distance to the powder convergence point as was shown in experiment 3. Depending on the deposited geometry desired, the shielding gas nozzle length can be increased, the mass flow can be increased, or the standoff distance can be decreased to increase powder catchment. These1096strategies are not dependent on the blown powder deposition head manufacturer, but rather can be utilized with any type of blown powder DED deposition head.AcknowledgementsThe authors would like to acknowledge the cooperation and support of Open Mind Technologies USA Inc. and the Mazak Corporation. This work is funded by the Department of Energy DE-EE0008303 with the support of Oak Ridge National Laboratory.

and the Mazak Corporation. This work is funded by the Department of Energy DE-EE0008303 with the support of Oak Ridge National Laboratory. This manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy. Research was co-sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office.References[1] Strong, D., Kay, M., Conner, B., Wakefield, T., and Manogharan, G., 2018, “HybridManufacturing – Integrating Traditional Manufacturers with Additive Manufacturing (AM) Supply Chain,” Addit. Manuf., 21, pp. 159–173.[2] Frazier, W. E., 2014, “Metal Additive Manufacturing: A Review,” J. Mater. Eng. Perform.,23(6), pp. 1917–1928.[3] Oyelola, O., Crawforth, P., M’Saoubi, R., and Clare, A. T., 2016, “Machining of Additively Manufactured Parts: Implications for Surface Integrity,” Procedia CIRP, 45, pp. 119–122.

pact Of Polyurethane Droplets on a Rigid Surface for Ink-Jet PrintingWenchao Zhou, Drew Loney, Andrei G. Fedorov, F. Levent Degertekin, David W. Rosen The George W. Woodruff School of Mechanical Engineering Georgia Institute of Technology 813 Ferst Drive, Atlanta, GA, 30332-0405Abstract Ink-jet printing enables a more efficient, economic, scalable manufacturing process for a wider variety of materials than other traditional additive techniques. Understanding droplet impact onto a substrate is critical for accuracy control and optimization of the droplet deposition process. This paper reviews fundamentals of droplet impact behavior from literature, followed by our CFD simulations of droplet impact on a substrate. A numerical model is developed to simulate the transient flow dynamics during impact and spreading of an impinging droplet on a wetting rigid wall. The phase-field surface tracking technique is employed within a fixed (Eulerian) structured mesh.

The phase-field surface tracking technique is employed within a fixed (Eulerian) structured mesh. Simulation results are critically compared and validated with predictions of analytical models and experiments from literature. Droplet impact and spreading of a thermoplastic polyurethane material are investigated for SFF applications.1. IntroductionInk-jet printing manufacturing is an emerging technology which employs printing of a liquid- based precursor in an additive manufacturing process to create three-dimensional products. Streams of droplets, such as molten engineering materials, suspensions, or photopolymer liquids are ejected onto a solid substrate to build up intricate components by maneuvering both the print head and substrate. Although ink-jet printing manufacturing is conceptually an elegant technology, in practice it has not proven easy to implement because of a number of difficulties.

Although ink-jet printing manufacturing is conceptually an elegant technology, in practice it has not proven easy to implement because of a number of difficulties. These include control of the final shape of droplets, splash control, failure of droplets bonding to each other, formation of pores at the interface between droplets, and residual stress in the final part, among others. Achieving a good quality of the final parts requires accurate control of the whole deposition process, which demands a deep understanding of underlying physics for each deposition step.Although the interaction of drops with surfaces has been extensively investigated for over a century, the development of comprehensive, predictive models has been difficult due to the complexity of the process and the interactions among many physical phenomena, including fluid mechanics, heat transfer, phase change, and surface chemistry.

This paper first reviews previous research on this subject and identifies the key issues involved in understanding droplet impact behavior. Then, a numerical model is presented to help understand droplet impact using the phase-field surface tracking technique. Simulation results are critically compared with analytical models and experiments from the literature. Upon validation, the model is used to study the impact behavior of polyurethane droplet, a material of interest which has potential SFF applications.2. Literature ReviewTo gain a better understanding of the droplet deposition process, previous research has been investigated, including droplet impact, droplet hardening and the coupling between them.Droplet impact may result in various outcomes depending on the circumstances of impact. Upon impact on a solid surface, droplets can stick to the surface, bounce off the surface, spread on the surface or splash under different impact conditions [1].

Upon impact on a solid surface, droplets can stick to the surface, bounce off the surface, spread on the surface or splash under different impact conditions [1]. Often the conditions leading to spreading and splashing are of particular interest. Due to the complexity of large deformation of the free surface and boundary conditions, different methods are employed to investigate this phenomenon, including experimental study, theoretical analysis and computer simulation.When a droplet comes out of the ejector, it travels with the initial ejection velocity under the sum of gravity force, lift force and drag force caused by the surrounding atmosphere. Different models have been developed to approximate the impact velocity for the cases of high Reynolds number and low Reynolds number [2].

Different models have been developed to approximate the impact velocity for the cases of high Reynolds number and low Reynolds number [2]. Droplet oscillations during flight were also modeled based on the droplet size [3].When the droplet reaches the substrate, different observations have been made for different wettability cases between liquids and solids and the concept of contact angle was introduced to quantify this phenomenon. Contact angle hysteresis [4] is also observed when the droplet spreads out on the substrate, which refers to the difference between advancing contact angle and receding contact angle. It is believed that this difference is caused by surface inhomogeneity, surface roughness, impurities on the surface and temperature [5-6].

It is believed that this difference is caused by surface inhomogeneity, surface roughness, impurities on the surface and temperature [5-6]. Due to the complexity of droplet impact behavior caused by the dynamic variation of the contact angle, experimental studies have been performed which suggest that the time evolution of the impact process can be divided into four distinct phases: the kinematic phase, the spreading phase, the relaxation phase and the wetting equilibrium phase [7]. These four impact regimes are classified according to the combination of driving forces and resisting forces to simplify the problem [8].Increasing the impact velocity sufficiently will lead to a splash [9], which is undesirable for SFF. Three different kinds of splash have been identified according to their reasons, which are surrounding gas [10], surface roughness [11] and compliance of the solid substrate [12].

Three different kinds of splash have been identified according to their reasons, which are surrounding gas [10], surface roughness [11] and compliance of the solid substrate [12]. The splash caused by surrounding gas is referred as corona splash, while the splash caused by surface roughness and solid substrate compliance is referred as prompt splash. Previous experiments have shown that the splashing threshold follows the empirical relation WeRe1/2=K, where K is a constant that depends on external parameters, such as surrounding gas pressure, surface roughness, etc. (Here, We refers to Weber number and Re refers to Reynolds number).Theoretical analysis has been performed to help understand the impact process, including dimensionless parameters that allow generalization and construction of regime maps. An energy conservation analysis can be used to obtain a good solution of the final spreading radius of the droplet after impact with Reynolds number between 150 and 2000.

An energy conservation analysis can be used to obtain a good solution of the final spreading radius of the droplet after impact with Reynolds number between 150 and 2000. Scale analysis can be used to get an estimation of the spreading velocity and time scales of driving and resisting forces under specific impact conditions [13]. To completely describe the whole impact process, 3-D differential equations are required to be built based upon a complete physical analysis, which is, however, far from being achieved. Therefore, mostly simplified 1-D models have been applied by only considering one parameter (the wetted area or the height of the droplet) to analyze the impact dynamics. Several 1-D models have been reviewed along with the dynamic contact angle models they used [14].Numerical modeling has been also performed for the prediction of droplet impact behavior.

Several 1-D models have been reviewed along with the dynamic contact angle models they used [14].Numerical modeling has been also performed for the prediction of droplet impact behavior. Various numerical methods have been investigated, including Marker-and-Cell (MAC), Volume-Of-Fluid (VOF), Level-Set, Phase-Field, and moving grids methods to predict droplet interface evolution upon transport towards and interaction with the substrate [15-19].To build a 3-D structure, droplets must solidify after impact in a predictable manner. Three different approaches to solidification have been investigated in this field: droplets freezing, solvent evaporation (e.g., water) if a suspension solution is used, and photopolymer curing upon exposure to light or radiation.

In the case of droplet freezing, several theoretical models for solidification of droplets have been reviewed, which suggest that the assumption that the droplet spreads first and then cools and solidifies is only valid within a limited parametric domain [15]. Most of the time, the spreading and solidification phenomena occur concurrently and are intimately coupled. Experimental studies on solidification, most of which use metal or hydrocarbon droplets of (1-4) mm in diameter with velocity of (1-3) m/s for easy observation [20], have also been undertaken and compared to numerical modeling, including development of different methods for treating the latent heat associated with phase change.Drying the droplets containing suspended or dissolved solids is another way to obtain solids [21]. Solvent evaporation regimes have been identified, and different models have been developed with different assumptions [22]. Complimentary experimental studies were performed, as well [23-24].

Solvent evaporation regimes have been identified, and different models have been developed with different assumptions [22]. Complimentary experimental studies were performed, as well [23-24]. Curing of photopolymer is another way to make a liquid droplet become solid [25]. Several mature industry applications of this method have been developed, including rapid prototyping machines from 3D Systems and Objet Technologies [26].3. Numerical modeling Since droplet impact could be very complicated in the most general case, several assumptions are made to simplify the analysis for the impact conditions of interest. a) Since the impact velocities of interest are around 10m/s, which is much less than the speed ofsound speeding the fluid, the fluid is assumed to be incompressible.b) The shape of the droplet may oscillate from prolate to oblate during fall because of the effect of the surrounding atmosphere, which is negligible for droplet diameters less than about 3 mm [27].

Since we are concerned with much smaller (10s of micrometer in diameter) droplets, it is assumed to be spherical throughout the transit period and at impact.c) The fluid flow is assumed to be Newtonian, and both liquid and gas flow are treated as laminar based on magnitude of the Reynolds number. The viscosity and surface tension are assumed to be constant during the impact.d) The droplet and substrate are assumed to be isothermal.With the assumptions above, the equations of conservation of momentum and mass can be formulated as follows:(1)(2)(cid:1)(cid:2)(cid:3) (cid:2)(cid:4)(cid:15) − ∇ ∙ [−(cid:9)(cid:10) + (cid:12)(∇(cid:3) + (∇(cid:3)))] + (cid:1)(cid:3) ∙ ∇(cid:3) = (cid:18)where: (Pa);(cid:12)∇ ∙ (cid:3) = 0is the fluid's density (kg/m3); u is the velocity vector (m/s); p represents the pressure s); F is the body force and I is the identity matrix.

∙(cid:1) denotes the dynamic viscosity (Pa(1)(2)Instead of directly tracking the interface between the liquid and gas, as well as the interface between liquid and solid using a moving (Lagrangian) grid, an Eulerian approach is used to locate the position of the interfaces on a fixed grid using the Phase-Field method.3.1. Phase-Field Method Different techniques exist for tracking an interface evolution numerically. The level-set method simply convects the interface with flow field, whereas the volume-of-fluid (VOF) method ensures volume conservation while convecting the interface. In this paper, the interfaces are tracked by the Phase-Field method, which not only convects the fluid interface but also ensures that the total energy of the system diminishes correctly. as described in [19]. A brief description of this method is given below.In the Phase-Field method, the interface is diffuse and is defined by a dimensionless parameter ϕ that varies from -1 to 1.

=(cid:22)(cid:25) (cid:24) (cid:23)3.2. Numerical model The numerical model is implemented using the commercially available software COMSOL 3.5a. A typical case is shown here to illustrate the implementation of the numerical model. Due to problem symmetry, the droplet and substrate can be modeled as shown in Figure 1, where the droplet comes out of the ejector at a certain distance from the substrate with an initial velocity. The droplet travels downward due to inertia and the gravity force and impinges on the rigid wall and spreads.DropletOpen boundaryInitial fluid interfaceSurrounding airAxis-symmetric axisRigid wallFigure 1 Geometry and boundary conditionThe initial configuration is shown in Figure 1 and the diameter and initial velocity of droplet are 50 µm and 10m/s (downward), respectively. The boundary conditions are also shown in Figure 1 and the equilibrium contact angle between the droplet and the substrate is set to π/2.

The boundary conditions are also shown in Figure 1 and the equilibrium contact angle between the droplet and the substrate is set to π/2. Table 1 gives the properties of droplet fluid and air used in the model simulations.Table 1 Properties of materialsMedium Density Droplet AirDynamic Viscosity Surface Tension1,000 kg/m3 0.01 Pa·s 1.225 kg/m3 1.789e-5 Pa·s0.07 N/mSince the phase field method is not intrinsically mass conserving, the ideal situation would be of no artificial mass loss during the simulation, which would indicate high accuracy of simulations. Figure 2 shows the evolution of the droplet mass as function of time, as it moves from its initial(10)position towards the substrate.

Figure 2 shows the evolution of the droplet mass as function of time, as it moves from its initial(10)position towards the substrate. The maximum mass loss of the droplet during the entire simulated period is around 3%, and it can be reduced by further refining the mesh.The form of the droplet during impact is often topologically-complex, but for SFF application the key factor is the spreading radius, which defines the resolution of this fabrication technique. Spreading radius is defined as the radius of the wetted area of the substrate. The change of spreading radius with time is shown in Figure 3, where we can see the spreading (up to about 15 micro seconds) and receding (beyond 15 micro seconds) of the droplet until it eventually reaches its equilibrium shape at about 150 micro second into the process.Figure 2 Mass change of the droplet due to numerical approximation of the interfaceFigure 3 Change of spreading radius with time during the droplet impact4.

Validation In order to validate the numerical model developed, a study of grid independence is conducted first, which includes assessment of the influence of the simulation domain extent and the mesh size refinement. The results of several static and dynamic analytical models are compared with those of the simulations. Finally, the simulation results are compared to the experimental results from literature for the impact condition of interest to SFF.4.1. Grid & domain independence The 'Open Boundary' condition is used in CAMSOL to simulate an infinite domain, and we examined if this boundary condition has any effects on the simulation results with different simulation domain sizes. To this end, the boundary length (the length of the boundary marked as 'open boundary' as Figure 1) was varied from 100µm to 300µm, and the simulation results showed a negligibly small change upon this variation.

Therefore, the boundary length of 100 µm is sufficient to obtain accurate results for fluid flow simulations.The grid independence validation ensures that the simulation results do not change significantly with different mesh sizes. In our numerical experiments, the minimum grid size of the mesh was changed from 4µm to 2µm and the results show that the mesh size of 4µm can provide sufficient accuracy in predicting the spreading radius and mass conservation.4.2. Analytical validation4.2.1. Equilibrium spread factor The spread factor is defined as the ratio of spreading radius to the initial radius of the droplet, and the equilibrium spread factor is the spread factor when the droplet stops spreading, which is very important to determine the final shape of the droplet. Two analytical models for the equilibrium spread factor are compared with our simulation results.

Two analytical models for the equilibrium spread factor are compared with our simulation results. The first model is the spherical model, which assumes that the droplet during spreading maintains a sphere cap shape [28].(11)789:;< = 2[@=>? (4 − 6DE=A8 + 2DE=A8@A)(/@ ]The simulation gives the equilibrium spread factor of βse=1.2160, while the analytical result is βequil=1.2599, yielding the relative error |βequil- βse|/βse=3.61%. The second analytical model is cylindrical model, which assumes the droplet spreads as a self-similar cylinder [28].(12)789:;< = 2[1 (6(1 − DE=A8)(/@ ]The equilibrium spread factor for this model is βequil=1.1006, and the relative error as compared to simulations is s |βequil- βse|/ βse=.9.49%.It can be seen that the simulation result is between that of the spherical and cylindrical models, where two extreme approximations are taken.

The droplet shape under these impact conditions is neither a sphere cap nor a cylinder, but is more close to a sphere, which agrees well with results in the literature. Specifically, results demonstrate that, for viscous fluids impacting at low velocity, droplet shapes are close to spherical for the entire deposition process, while droplets of(11)(12)low viscosity fluids have a spherical shape after reaching equilibrium, but not during dynamic spreading [28].4.2.2. Maximum spread factor Maximum spread factor is defined as the ratio of the maximum spread radius of the droplet to its initial radius. The maximum spread factor dictates how the spacing between nozzles in the array needs to be arranged and how the neighbor droplets interact with each other.

The maximum spread factor dictates how the spacing between nozzles in the array needs to be arranged and how the neighbor droplets interact with each other. There are many empirical [29-30] or semi-empirical [31-32] models to predict the maximum spread factor and the following empirical equation was shown to provide better predictions than other models [33]:(13)F0.2HℎJ.@@KLJ.MMN+1 4(cid:24)(1 − DE=A8)O 7PQR+2 3S( 7PQR=KL 12+ 1where Oh is the Ohnesorge number, We is the Weber number, βmax is the maximum spread factor, and θe is the equilibrium contact angle. Two simulations were run under different impact conditions and compared with the analytical model (Eq. 13), as shown in Table 2.

Two simulations were run under different impact conditions and compared with the analytical model (Eq. 13), as shown in Table 2. It can be seen the simulation results agree well with the analytical results.Table 2 Comparison of maximum spread factor between simulations and analytical modelDroplet RadiusImpact VelocityDynamic ViscositySurface TensionDroplet DensitySimulation ResultsTheoretical ResultsCase125µm10m/s0.01Pa·s0.07N/m1e3kg/m31.76321.7476Case230µm15m/s8.94e-4Pa·s0.072N/m1e3kg/m33.066734.2.3. Dynamic spreading In this section, the results of three well known 1D models for dynamic evolution of the wetted area/height of the droplet are compared with those of our numerical model under certain impact condition. The main challenge of the 1D model lies in the assumed physics of viscous dissipation and in the lack of self-consistency between the description of kinetic energy and that of dissipation. Bechtel et al.

The comparison of results between the simulation and BBT model is shown in Figure 4, from which we can see that the results agree very well when the coefficient-multiplier of γ goes to 40. It is noteworthy that there is no rigorous calculation of γ and only a rough estimation can be made here.Figure 4 Comparison between simulation and BBT modelAfter that, various 1D models have been proposed [14, 29-33, 35]. Since the free surface of low viscosity droplets at its first maximum spreading is far from being spherical, the cylinder form was adopted by Kim et al. [35], which is called the KC model:1 12T T(cid:4)kY1 2+1 271 hM] YTh T(cid:4)](cid:24)m +T T(cid:4)Fh(cid:24)(1 − DE=A8) +1 3hO +(cid:21) 2Hℎ h(cid:24)YTh T(cid:4)(cid:24)]= 0with the initial conditions:S(/(cid:24)( @ng = 1; and the coefficientcn c` |`aJ = √KL o( (cid:24) +(( nqr(cid:24) hJ(1 − DE=A8) +(cid:24)pwherein the parameter γ is adjustable.

The comparison is shown in Figure 6, which indicates that when the coefficient-multiplier goes to 0.45, the early stage of the spreading agrees well, and the long-term spreading dynamics predictions agree well between each other when s takes its default value.Th T(cid:4)2 31 hM]|`aJ = √KL Y+wherei fg ; = = 1.41Hℎh =](cid:24)(16)Figure 6 Comparison between simulation and AGM model5. Experimental validationTo further validate the numerical model, experimental validation is necessary. Although there is a large body of experimental results in literature, few of them focus on the droplet size at the micron scale due to the difficulty of observation. In order to obtain similar droplet impact behavior with our interests, an experiment with a Weber number of interest was chosen for comparison.

In order to obtain similar droplet impact behavior with our interests, an experiment with a Weber number of interest was chosen for comparison. A water drop impact on a steel substrate with Weber number of 27 and contact angle of π/2 [31] is selected to compare with the simulation results, as shown in Figure 7, which is quite excellent during the early stage of the spreading (left panel).Figure 7 Comparison between experimental data and simulation: early stage(left); all time (right)6. Impact of Polyurethane droplet A thermoplastic polyurethane material is of particular interest in this paper due to its potential SFF applications. The fluid properties of polyurethane are taken from [36] as shown in Table 3.

The fluid properties of polyurethane are taken from [36] as shown in Table 3. Table 3 Impact conditions and fluid properties of polyurethaneDroplet RadiusContact AngleImpact VelocityDynamic ViscositySurface TensionDroplet DensityValue25µmπ/210 m/s3 Pa·s0.029 N/m1050 kg/m3Using the simulation conditions from Table 3, simulation results are shown in Figure 8. The time scale of spreading is increased, as compared to water, to about 0.1 second due to the higher viscosity of polyurethane and the maximum radius of the wetted area is only slightly larger than the equilibrium radius with no oscillation.

As such, droplet impact behavior for high viscosity fluids is simpler than for low viscosity fluids, indicating that manufacturing process control may be simpler than in current printing processes.Figure 8 Simulation result of the spreading of polyurethane dropletThe BBT model and KC model are fundamentally unable to predict the spreading of the fluid at such high viscosity because they do not incorporate all the physics of viscous dissipation, as in the AGM model. Simulation results are compared with that of the AGM model as shown in Figure 9. The AGM model predictions compare well with simulations on the time scale for droplet spreading before reaching an equilibrium state. However, there are qualitative differences in dynamics of the droplet behavior (e.g., the AGM model does not predict a droplet recoil), as well as the quantitative discrepancy in the maximum and final spreading radii.

Yes, the AGM model may be sufficiently accurate to make quick estimations for the process design, while using detailed and computationally intensive simulations for process refinement and optimization.Figure 9 Comparison between simulation and AGM model7. Conclusion Previous research on droplet impact and hardening has been reviewed for SFF using ink-jet printing methods. A numerical model has been developed to simulate the transient fluid dynamics during droplet impact and spreading using phase-field method within an Eulerian fixed grid. A typical case under impact conditions of interest is presented to illustrate the implementation of the numerical simulation. Then two critical parameters for manufacturing, namely the equilibrium and maximum spreading radii, are calculated using both validated empirical analytical models and our numerical model and the comparison shows that the results agree very well.

Further, three well known 1-D models are then used to calculate the dynamic spreading process which is compared with the spreading process computed by the proposed numerical model. By tuning the adjustable parameters of the analytical models, good agreements between the simulations and analytical results has been demonstrated. Complimentary to analytical validation of the numerical model, experimental data from literature was successfully used to validate the numerical model under relevant impact conditions. Finally, the impact of polyurethane droplet that is of particular interest for potential SFF applications is simulated using the proposed numerical model and compared with those three 1-D models. Results illustrate that the detailed understanding of the spreading process, as enabled by the CFD simulations, is essential to predict droplet behavior for ink-jet printing of complex engineering materials.

Further, results show that only one of the simplified 1-D models (the AGM model) is able to predict the impact behavior of high viscosity fluids due to its incorporation, albeit rather simplistic, of viscous dissipation. AcknowledgementWe gratefully acknowledge the U.S. National Science Foundation, through award DMI-0900322. Any opinions, findings, and conclusions or recommendations expressed in this publication are those of the authors and do not necessarily reflect the views of the National Science Foundation.References[1][2] [3][4][5][6][7][8][9][10][11][12][13][14]A. L. Yarin, "Drop impact dynamics: Splashing, spreading, receding, bouncing," Ann. Rev. Fluid Mech, vol. 38, pp. 159-192, 2006. K. Mramor, "No Splash on the Moon," University of Ljubljana2007. "Deformation S. http://www.shorstmeyer.com/wxfaqs/float/dropdeform.html, ed, 2008. E. B.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIMPACT OF POROSITY TYPE ON MICROSTRUCTURE AND MECHANICAL PROPERTIES IN SELECTIVELY LASER MELTED IN718 LATTICE STRUCTURESS. Ramachandra*, B. B. Ravichander*, B. Farhang*, A. Ganesh-Ram*, M. Hanumantha*, J. Marquez*, G. Humphrey*, N. Swails+, A. Amerinatanzi*†Mechanical & Aerospace Engineering, University of Texas at Arlington, Arlington, TX. + Math, Science, Nursing, and Public Health, University of South Carolina Lancaster, SC. †Materials Science and Engineering, University of Texas at Arlington, Arlington, TX.AbstractLaser Powder Bed Fusion (LPBF), one of the most employed additive manufacturing techniques for metals, has opened new dimensions in realizing strong and weight reducing structures. In this study, Inconel 718 (IN718) unit cell designs, were fabricated through the LPBF technique and analyzed.

In this study, Inconel 718 (IN718) unit cell designs, were fabricated through the LPBF technique and analyzed. Among the plethora of lattice structures in existence, BCC, BCC-Z, FCC, FCC-Z, Gyroid, Diamond and Schwarz structures have been selected to focus on. A relationship between the mechanical properties including yield strength, failure stress and strain, and hardness with each type of unit cell was established. Also, the effect of the possible defects on the hardness value was examined using microstructural analysis on samples. Scanning Electron Microscopy (SEM) analysis was also performed to examine the possible defects and its effect on the hardness of the as-built part. The SEM images of the grain structures indicated higher levels of isotropy in Gyroid, and Diamond samples compared to the rest of the samples which relates to the load bearing capacities of each unit cell structure.

A similar trend was observed in terms of the uniformity of meltpool which can be linked with the consistency in yield characteristics. Further, Diamond and BCC-Z structures displayed high values of hardness in comparison with rest of the samples.Keywords: Laser Powder Bed Fusion, Porous Structures, Mechanical properties, IN718, Microstructure, Vickers Hardness.1. IntroductionThe Inconel 718 (i.e., IN718), a superalloy with nickel as its base constituent, demonstrates extraordinary properties such as good impact resistance, excellent tensile and compressive strengths [1, 2]. It also possesses high corrosive resistance on account of elevated levels of nickel (estimated to be between 50 – 55 %) [3, 4]. IN718earns the classification of a superalloy due to its ability to retain the chemical, mechanical, and metallurgical properties at elevated temperatures of the range 650 to 725 degree Celsius [5].

IN718earns the classification of a superalloy due to its ability to retain the chemical, mechanical, and metallurgical properties at elevated temperatures of the range 650 to 725 degree Celsius [5]. With such excellent set of properties offered, the superalloy has been used in aerospace, automobile, and military applications [6-11].The rapid advancements in technology have created a great necessity for light weight and complex structures. Implementing engineered porous and Triply Periodic Minimal Surface (TPMS) based structures into part designs provides a feasible opportunity to greatly reduce the weight of the parts. However, conventional modes of manufacturing limits the ability to produce these complex structures and hence Additive Manufacturing (AM) techniques can be employed to1000accomplish the same [12].

Among various metal AM techniques at disposal, Laser Powder Bed Fusion (LPBF) is one of the widely used methods that can achieve fully dense metal parts with complex geometrical shapes [12-15]. Unlike conventional fabrication methods, the LPBF technique is capable of manufacturing IN718 parts with engineered porosity embedded into their design Due to the excellent weldability of IN718, complex lattice structures can be fabricated using the LPBF process.Distinct types of porosity shapes can be implemented into the designs such as BCC, BCC- Z, FCC, FCC-Z, Honeycomb and Gyroid. The unit cell shapes tend to have a direct impact on the mechanical attributes of the fabricated parts. While some unit cell types display excellent compressive capabilities, other unit cell types may lack the compressive capability meanwhile display good hardness or tensile strength.

While some unit cell types display excellent compressive capabilities, other unit cell types may lack the compressive capability meanwhile display good hardness or tensile strength. A study conducted by Leary et al., in which the lattice performance of FCC, BCC, FCC-Z, and BCC-Z was investigated under the uniaxial compression, a large collapse stress to plateau stress ratio was observed on the different models [16]. Similarly, Maskery et al. [17] performed uniaxial compression tests on Al-Si10-Mg BBC structures and found approximately 95% of the structure loss occurred at 9% strain rate. Also, the microstructure of the parts can be affected by the unit cell types which thereby lead to the variation in mechanical properties.

Also, the microstructure of the parts can be affected by the unit cell types which thereby lead to the variation in mechanical properties. To obtain the desired mechanical characteristics, it is crucial to establish a relationship between the cell types and their effect on the various mechanical properties.Several studies have been performed to investigate the effect of porous structures on the mechanical properties. Sing et al. performed compressive tests on Ti6A14V using different cellular structures and showed the effect of the process parameters on the mechanical properties as well as the elasticity of lattice structures [18]. It was concluded that at an increase of 50% in strain would render the structure incapable of regaining its original strength [19]. Qiu et al. performed a study to find how the laser power influences a diamond lattice of AlSi10Mg [5]. The samples were made using 150W and 400W laser powers.

Qiu et al. performed a study to find how the laser power influences a diamond lattice of AlSi10Mg [5]. The samples were made using 150W and 400W laser powers. It was concluded that there is a linear relationship between the power and the strut diameter. In addition, they found an increase in melted powder sticking to the surface of the strut as the laser power increased [20].Despite the aforementioned studies on the lattice structures, the effect of unit cell shapes on the mechanical properties and the microstructure have not been extensively researched. In our previous study, uniaxial compression tests were performed on IN718 porous structures to find the effect of cell shape on the mechanical properties. Also, microstructural analysis was performed to comprehend the effect of unit cell types on parameters such as yield and failure stresses. Among all the structures, FCC-Z showed the highest yield strength followed by BCC-Z.

Among all the structures, FCC-Z showed the highest yield strength followed by BCC-Z. Similarly, in TPMS structures, Schwarz showed the highest yield strength but was significantly lower in comparison with the BCC-Z and FCC-Z structures. Following a similar trend, the melt pools of TPMS structures, especially in Gyroid and diamond were uniform in terms of depth and width whereas irregularities were observed in the other samples. This can be linked to the nature of their corresponding yield values. Upon observing the SEM images of the samples before and after compression tests, BCC, FCC, and Schwarz samples retained the shape of their unit cells to a recognizable extent post loading. However, maximum displacement was evident in Gyroid followed by Schwarz and the Z-strut counterparts outperformed the BCC and FCC porosity types [21].10012.

However, maximum displacement was evident in Gyroid followed by Schwarz and the Z-strut counterparts outperformed the BCC and FCC porosity types [21].10012. Materials and methods2.1 Materials and FabricationThe IN718 powder used in this study was obtained from EOS (EOS GmbH, Krailling, Germany) and had the composition of Nickel (55%), Chromium (17%), Niobium (4.75%), Molybdenum (3.3%), and the remaining weight percentage consisted of aluminum, cobalt, and phosphorous. The average size of the powder particle was determined to be around 40 microns. An EOS M290 metal 3D printer (EOS GmbH, Electro Optical Systems, Krailling, Germany) equipped with a 400W Ytterbium fiber laser used the aforementioned powder for LPBF fabrication of the samples.The part design process involved the creation of geometries for each of the BCC, BCC-Z, FCC, FCC-Z, Gyroid, Schwarz and Diamond lattice structures in the form of a cube with the dimensions 8 x 8 x 8mm.

SolidWorks (Dassault Systemes, Vélizy-Villacoublay, France) was used to create the necessary CAD files. Though all the unit cells had the same dimensions of 2 x 2 x 2 mm, the TPMS structures, namely the Gyroid and Schwarz unit cell designs were conceived with the help of K3Dsurf and Materialise Magics (Materialise 2020, Belgium).Figure 1. CAD models of a) BCC, b) BCC-Z, c) FCC, d) FCC-Z, e) Gyroid, f) Schwarz, and g) Diamond2.2. Experiment MethodThe samples were tested for hardness, mechanical properties, and microstructural analysis. The hardness values were determined on the Metal-tester model 900-391D (LECO, Missouri, USA). A Shimadzu EHF 100 fatigue testing machine was used to carry out uniaxial compression1002test. The microstructural analysis involved the use of a Scanning Electron Microscope (SEM), Hitachi S-3000N Variable Pressure SEM (Hitachi, Canada).3. Results and Discussion3.1.

The microstructural analysis involved the use of a Scanning Electron Microscope (SEM), Hitachi S-3000N Variable Pressure SEM (Hitachi, Canada).3. Results and Discussion3.1. Uniaxial Compression TestFigure 1 shows the compression test results for the samples with different lattice structures [22, 23]. As can be seen in figure 1, though Gyroid and Diamond showed excellent compressive displacements of 4.836mm and 4.083mm respectively, they lacked in yield stresses. Since the maximum force used for the compression was 83KN, the overall displacement of these two specimens were not recorded to the full extent. However, this prolongation in failure can be associated to the smoothness in the transition of geometry of struts in each unit cell. The Z-variants of both BCC and FCC showed a familiar trend of outperforming their simple counterparts which can again be credited to the presence of Z-struts.

The Z-variants of both BCC and FCC showed a familiar trend of outperforming their simple counterparts which can again be credited to the presence of Z-struts. Moreover, among all the samples including the TPMS structures, BCC-Z and FCC-Z attained the highest numbers in terms of compressive forces - 946MPa and 965MPa, respectively.8000070000)N60000BCC(e c r o fe v i s s e r p m o C50000400003000020000BCC-ZFCCFCC-ZGyroidSchwartz10000Diamond0013 Compressive displacement (mm)2456Figure 2. Compressive displacements and compressive forces sustained by each unit cell depicted graphically.3.2. Microstructure AnalysisThe Microstructural analysis was carried out using the SEM at a constant rate of magnification to reveal the nature of melt pools and the defects.

Microstructure AnalysisThe Microstructural analysis was carried out using the SEM at a constant rate of magnification to reveal the nature of melt pools and the defects. According to figure 2, while samples like BCC, BCC-Z, FCC, FCC-Z, and Schwarz displayed random irregularities in the widths and depths of adjacent melt pools, Diamond and Gyroid exhibited uniform spread of melt pool sizes. This nature of variations in melt pools can be credited to the transition in the geometry of the unit cells. In the case of BCC, BCC-Z, FCC, FCC-Z and Schwarz the transition in the shape of the unit cells is instantaneous leading to a lot of variations in the melt pools. Whereas, in Gyroid1003and diamond, the unit cells had a smoother geometrical transition which maintains similar and uniform development of melt pools.

Whereas, in Gyroid1003and diamond, the unit cells had a smoother geometrical transition which maintains similar and uniform development of melt pools. Uniformity in the melt pools ensures consistent and good yield results while irregularity diminishes the yield results.Further, some of the defects observed in the SEM images involves cracks, unmelted powder, precipitates, and pores. Some of the defects can be seen enclosed in red in figure 2. Following a similar trend as the melt pools, BCC, BCC-Z, FCC, FCC-Z, and Schwarz displayed notable amounts of partially and unmelted powder on the surface and this can again be attributed to the instantaneous transition in their geometry. Conversely, the amount of partially/unmelted powder observed for the diamond and gyroid samples were negligible. While the amounts of precipitates and pores were commonly at a miniscule range in all the samples, some cracks can be observed in FCC and BCC structures.

While the amounts of precipitates and pores were commonly at a miniscule range in all the samples, some cracks can be observed in FCC and BCC structures. These cracks tend to initiate premature part failures making them less reliable.1004Figure 3. SEM images of Melt pools and defects observed in a) BCC, b) BCCZ, c) FCC, d) FCCZ, e) Gyroid, f) Diamond, g) Schwartz samples.10053.3. Grain StructureThe key to achieve desired mechanical attributes of a part is in the manipulation of isotropy in grain structure. SEM images of grain structure for each unit cell type can be observed below in figure 3. Samples of BCC, BCC-Z, FCC, FCC-Z and Schwarz displayed significant levels of anisotropy in grain sizes and shapes. However, Gyroid and Diamond displayed excellent isotropic grain structure. It can be noted that, because of the geometric nature of its unit cell, the Gyroid sample showed the finest quality of isotropy among all the samples.

It can be noted that, because of the geometric nature of its unit cell, the Gyroid sample showed the finest quality of isotropy among all the samples. By analyzing the trends evident in the microstructural images of the grain structures, one can infer those higher levels of anisotropy can lead to early failures or reduction in the overall compressive displacement of the material.1006Figure 4. Grain structures of a) BCC, b) BCCZ, c) FCC, d) FCCZ, e) Gyroid, f) Diamond, g) Schwartz samples observed through SEM.10073.4. Hardness AnalysisThe data (figure 4) extracted from the Vickers hardness tests revealed that the Z-strut variants of both BCC and FCC outperformed their standard counterparts. Comparing all the structures, Diamond exhibited the greatest magnitude of hardness of 421 HV and Gyroid bearing the least hardness of 337 HV.

Comparing all the structures, Diamond exhibited the greatest magnitude of hardness of 421 HV and Gyroid bearing the least hardness of 337 HV. Apart from the Z-variants showing greater hardness than their simpler forms (BCC, and FCC), there is no trend evident in the variation of hardness values with respect to the unit cell shapes.Figure 5. Vickers Hardness values of each unit cell type.4. ConclusionThis study investigated the effect of type and shape of lattice structures on the mechanical properties, microstructure, and hardness of samples. It was revealed that unit cell geometry has a direct influence on the microstructural aspects including grain structure, defects, and melt pools successively impacting the mechanical properties of a part. It was found that unit cells with smoother transitions, displayed isotropic and defect-lacking microstructural quality, hence offering optimal mechanical properties.

It was found that unit cells with smoother transitions, displayed isotropic and defect-lacking microstructural quality, hence offering optimal mechanical properties. Additionally, a variation in the hardness value was observed among samples with different lattice types, with the highest and lowest magnitude of hardness being found for Diamond and Gyroid structures, respectively.5. References1.Li, J., et al., Microstructural evolution and mechanical properties of IN718 alloy fabricated by selective laser melting following different heat treatments. Journal of Alloys and Compounds, 2019. 772: p. 861-870.10082.3.4.5.6.7.8.9.10.11.12.13.14.15.Calandri, M., et al., Texture and Microstructural Features at Different Length Scales in Inconel 718 Produced by Selective Laser Melting. LID - 10.3390/ma12081293 [doi] LID - 1293. (1996-1944 (Print)). Wang, Z., et al., The microstructure and mechanical properties of deposited-IN718 by selective laser melting.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceIMPACT OF RAW MATERIAL COMMODITIZATION ON SUPPLY CHAINSS. H. Khajavi1 , G. Deng1, J. Holmström1, P. Puukko21 Department of Industrial Engineering and Management, Aalto University, Espoo, Finland 2 VTT research organization, Espoo, FinlandAbstractIn this research we analyze the impact of a less cost intensive raw material supply chain on the utilization of additive manufacturing (AM) and its implications for the whole production supply chain. Utilizing the case data and scenario modeling we compared the competitiveness of additive manufacturing in comparison to conventional methods such as machining and casting. Results illustrated the increased production competitiveness with regard to both conventional methods.

Results illustrated the increased production competitiveness with regard to both conventional methods. Moreover, we found simplicity and swiftness of supply chain as the two significant changes that occur as the result of a new raw material supply. One of the major limitations of this research was due to the secrecy of the companies which are currently utilizing the technology and their reluctance to share their perceived competitive advantage for the analysis. Therefore, the data extracted from our own case simulations which then were analyzed to compare the supply chains. Future research should expand the number of cases and utilize more suitable AM machines.Keywords: Additive manufacturing, raw material commoditization, supply chain.IntroductionAdditive Manufacturing (AM), a novel production technology that recently started to be used for the production of metal final parts in larger volumes (GE), still struggles with a number of shortcomings [3].

The cost and availability of raw material is one of the issues especially for the metal parts production. This makes the utilization of AM for final parts in a supply chain, risky. However, the benefits gained from the use of three-dimensional printing in some cases outweigh the uncertainty and difficulties. For instance in case of GE fuel injectors for the jet engines, the parts consolidations and performance improvements due to complex geometries were the important selection criteria. Another beneficial aspect of implementing AM can be the possibility of production decentralization [3], which significantly reduces the delivery time while increasing supply chain flexibility. Higher level of availability for an AM grade powder with an economical cost can unlock these benefits for more applications.Metalysis, is a novel process of titanium raw material production which promises much cheaper raw material in powder form factor, can change the current supply chains [9].

The fact that Metlysis yields a raw material which undercuts current titanium purification methods (e.g.: Kroll) by a wide margin, makes it a candidate not only for AM but also conventional manufacturing (CM) methods such as casting and CNC machining. Therefore, it is essential to1 1230know the actual impact on the production cost as well as the supply chain effect. In this paper we aim to answer that question.The remainder of this paper is organized as follows: Section 2 presents a literature review; Section 3 explains the research methodology; and Section 4 presents the findings and results of our analysis.

This paper ends with conclusions summarizing the research outcomes, and suggestions for future investigation are provided.Literature reviewIn this section we briefly review the literature on the two main subjects of this paperwhich are additive manufacturing supply chain and Metalysis process.Additive manufacturing and its supply chainASTM Standard defines Additive Manufacturing (AM) as “the process of joining materials to make objects from 3D model data, usually layer upon layer, as opposed to subtractive manufacturing methodologies, such as traditional machining”. The literature for this technology is rife with synonyms, including additive processes, additive fabrication (AF), additive layer manufacturing (ALM), layered manufacturing, and solid freeform fabrication (SFF), 3D-printing (3DP) and rapid manufacturing (RM).In Figure 1, the AM process is explained.

Process begins with development of a three- dimensional CAD model of the object with all its details and dimensions. This CAD file is then converted to stl-file, which most AM-machines understand. Next the 3D file is sliced into very thin two-dimensional (2D) cross sections (layers) by a computer program. The machine produces each layer on top of the previous one, completing the object in question, utilizing different solidification methods of raw material in its production chamber. The process takes from a few hours to a few days to produce an object, depending on its size and needed accuracy. After the object has been removed from the machine and cleaned, it can be post-processed. The post- processing methods vary according to the AM process, but they can include removal of support material, infiltration of pores, surface finishing (peening, blasting or tumbling), annealing, and cleaning and machining [8].2 1231Fig.1.

Additive manufacturing processAM has long history in conceptual and functional prototyping but now the step toward directly manufacturing of final parts and products has been made. The methods are reaching the suitable precision and quality necessary to be used as final functional parts for special applications, such as in aerospace, defense and other highly-engineered systems [10].The reason AM technologies have been gathering momentum is certain benefits over the conventional manufacturing methods. According to Holmsröm et al. [5] the most important ones are toollessness, production for performance and potential for simpler supply chains. Additionally Khajavi et al. [3] suggested that using AM methods could reduce material waste (by as much as 90%).Figure 2 is a simplified illustration of production process flow while utilizing traditional manufacturing methods in contrast to the AM method.

Utilization of AM alleviates the supply chain complexity while reducing production steps and enables a far more flexible decentralized supply chain [3].Fig.2. Change from conventional process flow to AM process flow (Rommel and Fischer, 2013)3 1232Metalysis processMetalysis Process is a new method for purification of titanium which yields metal powder. As it is essentially one step process, it is possible to reduce cost of the production considerably perhaps even up to 75%. Titanium oxide (sintered rutile sand) serves as cathode and carbon as anode. Both anode and cathode are placed in a bath of calcium chlorine salts at 800°C-1000°C temperatures. Salts acts as an electrolyte allowing current (metal ions) to travel from cathode to anode. In anode, ions react with carbon and forms carbon dioxide gas. Meanwhile the cathode transforms into metal [2,9].Fig.3. Metalysis Process vs.

In anode, ions react with carbon and forms carbon dioxide gas. Meanwhile the cathode transforms into metal [2,9].Fig.3. Metalysis Process vs. Kroll Process [9]When compared with the conventional Kroll process for titanium metal extraction (Table 1), Metalysis process offers many significant advantages. Hereby, the conventional extraction process is known as Kroll process, which extracts titanium in a non-powder but spongy form from the minerals [7]. Because of the complicated, energy and time-consuming procedures, Kroll is an expensive pyro-metallurgical process to extract titanium [1].Table 1.

Because of the complicated, energy and time-consuming procedures, Kroll is an expensive pyro-metallurgical process to extract titanium [1].Table 1. Comparison of industrial titanium production processes with Metalysis process [6]ProcessKrollFeedstock TiCl 4Reductant MagnesiumBy-products MgCl 2Duration Batches, up to 7 daysProduct SpongeMetalysis (FFC)TiO 2Applied current (electrons)CO, CO 2Semi-continuous, 8-24 hoursPellets/powderLiterature gapThe current body of knowledge is mostly focused on the investigation of Metalysis and similar metal extraction methods from the technical perspective. This makes the analysis of novel ways of metal extraction methods on the choice of production method and supply chain implications of such choices, unexplored. We aim to fill the gap in the literature with regard to the possibilities of a much cheaper raw material for novel and superior supply chains.

We aim to fill the gap in the literature with regard to the possibilities of a much cheaper raw material for novel and superior supply chains. Our research question is as follows;What are the supply chain effects of an unconventional raw material supply?4 1233MethodThe methodology chosen for this study is modeling of actual cases for the competitiveness comparison of the production methods and building up on top of with supply chain interpretations. We utilized case data, expert opinion in addition to well referenced articles and books. The cost modeling lead us to the selection of more competitive scenarios and the shifting points were calculated based on the lower costs. Our cases are two parts used on the commercial airplanes, first a low pressure turbine blade (LPB) and second, a jet engine loading bracket (ELB).

Our cases are two parts used on the commercial airplanes, first a low pressure turbine blade (LPB) and second, a jet engine loading bracket (ELB). Both parts are manufactured with Ti6Al4V which is among the most common titanium alloys.For each case we calculated the production cost utilizing additive and conventional manufacturing methods while using raw material from Kroll or Metalysis process. As a result of that analysis we ended up with eight models and respective cost values for each.ResultsAs mentioned, Metalysis process can reduce the cost of extracting high-purity titanium powder by 75% compared to the Kroll process. Since in this study the cost of Kroll Ti6Al4V- ingot and -bar is 32€/kg, then the Metalysis Ti6Al4V-powder price is calculated to be 8€/kg.

Since in this study the cost of Kroll Ti6Al4V- ingot and -bar is 32€/kg, then the Metalysis Ti6Al4V-powder price is calculated to be 8€/kg. Moreover, because the low pressure turbine blade (LPB) is a casted part and jet engine loading bracket (ELB) is CNC machined while analyzing the conventional manufacturing (CM) methods costs are presented in two separate tables. Table 2 presents the cost competitiveness of manufacturing methods with different sources of raw material for LPB when the production batch is assumed to be 1000 parts. Moreover, the assumption in case of Metalysis powder use for the casting is that no cost saving is gained from the use of raw material in powder form.Table 2.

Moreover, the assumption in case of Metalysis powder use for the casting is that no cost saving is gained from the use of raw material in powder form.Table 2. Per part manufacturing cost comparison of low pressure turbine blade (LPB) under different circumstancesKrollMetalysisAM73.01€53.35€CM (casting)36.65€31.03€A cost reduction of about 37% per LPB parts is achieved for the additive manufacturing while the source of raw material becomes Metalysis powder. However, the cost of casting LPB is only reduced by 18%. All in all, on the per part basis for the production batch of 1000 parts, casting is cheaper than AM. The per part production cost of AM and CM crosses at 318 while Kroll raw material is in use. This number increases to 432 when Metalysis powder is in use.

The per part production cost of AM and CM crosses at 318 while Kroll raw material is in use. This number increases to 432 when Metalysis powder is in use. This is a promising improvement in the competitiveness of AM over CM (casting in this case).Table 3 presents the cost competitiveness of manufacturing methods with different sources of raw material for ELB. The assumption in the case of Metalysis utilization for machining process is that 10% additional cost is added to the raw material cost for the powder solidification before the process.5 1234Table 3. Per part manufacturing cost comparison of jet engine loading bracket (ELB) under different circumstancesKrollMetalysisAM180.49€108.81€CM (machining)115.49€49.45€A cost reduction of about 66% per ELB parts is achieved for the additive manufacturing while the source of raw material becomes Metalysis powder. However, the cost of machining an ELB is reduced by 134%.

However, the cost of machining an ELB is reduced by 134%. All in all, on the per part basis for the production batch of 1000 parts, machining is cheaper than AM. The per part production cost of AM and CM crosses at 6 while Kroll raw material is in use. This number merely increases to 7 when Metalysis powder is in use. This is a marginal improvement in the competitiveness of AM over CM (machining in this case).As we illustrated above, AM gains additional competitiveness while the raw material of choice (the more cost effective source) is also in powder form and this has impacts on the supply chain. The first most obvious effect is the reduced complexity of raw material supply chain and therefore higher reliability.

The first most obvious effect is the reduced complexity of raw material supply chain and therefore higher reliability. Metalysis as a single step method of extraction can contribute to the supply chain robustness and lower the supply interruptions due to market demand fluctuations.Decentralization of production is another potential supply chain improvement that increases flexibility and reduces excess costs related to inventory keeping and transportations [3]. If the value of the low lead time can justify the additional production cost (for instance in a maintenance setting), then AM can deliver the intended benefits in a decentralized setting.Another known benefit of AM is risk reduction in uncertain market settings [4]. In these cases (especially the LPB), improvements in AM switch over (to CM) volume enables a longer period of AM utilization and allows market failure risk reduction (for example in a new product introduction setting).

This can be extremely valuable as it allows the manufacturer to postpone the tool production and accelerates the supply chain.And if we take all the above mentioned factors into account, utilization of AM with a Metalysis like raw material provision method will potentially lead to simplification throughout supply chain. From shortening of the raw material production chain to the faster parts manufacturing (without long tool production lead times) and in-situ delivery. Therefore, the planning and control will be reduced and partly eliminated (no need for inventory and less sub- contractors and parts transshipments).ConclusionsIn this research we utilized real world cases to illustrate the impact of a phenomenon the cost (Lower cost, more reliable, Metalysis raw material extraction method) on competitiveness of additive manufacturing and how that can change supply chains.

Results showed that in both cases a much cheaper source of raw material powder will improve AM cost6 1235competitiveness over conventional methods; however the impact is of different magnitude for different cases and production methods. Finally we summarized the implication of such AM feasibility improvements on the supply chains.This study was performed using the available selective laser melting (SLM) metal machines which are meant for research. However since the economics of scale for additive manufacturing to some extend can result in more cost effective parts while the production chamber becomes larger, we think future research should address this issue. Moreover, we encourage the researchers to expand the results of this paper with replicating the case results on an electron beam melting (EBM) machine and update the supply chain implications based on the those results.References1. Qian, M. & Froes, F. H. (Eds.). (2015).

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperImpact of Solid Freeform Fabrication in enabling design and prototyping capabilities for Competitive Robotics such as World Robotics League, FTC, FRC, WRO etc.Rajeev Dwivedi*, a , Ravi Bhupathiraj,a, Rohit Bhupathirajub, Anvita Agarwal c, Indira Dwivedi d, Bharat Dwivedi d, Saketh Nadella ea: STEM and Robotics Academy, Plano TX, b: Allen ISD, Texas, c: Frisco ISD, TX, d: LWSD ISD, WA, e: Plano ISD, TX, Abstract:Competitive Robotics such as World Robotics League (WRL), World Robotics Olympiad (WRO), BEST, FIRST programs such as FTC and FRC, etc. have recently evolved as an avenue for learning, innovation and providing real life perspective to classroom learning. Given the short duration for Robot Development and diverse range of challenges, innovation, and product development has to happen at fast pace.

Given the short duration for Robot Development and diverse range of challenges, innovation, and product development has to happen at fast pace. Typical challenges include stow, configure, launch, traverse 3D environment, sorting, assembly etc. Many challenges include underwater and aerial spaces. Participants typically use reconfigurable “Robotics Kits”. However, "Robotics Kits" are limited in many ways and do not go beyond mobile platform and simple mechanisms. Many participants are turning to quicker alternatives such as Solid Freeform Fabrication. This paper provides a current state of the Competitive Robotics, the existing solutions, and examples of how Solid Freeform Fabrication is accelerating innovation in competitive robotics.1. IntroductionAbility to correlate the classroom learning with real life system, verification and qualification of ideas via experimentation is cornerstone of learning and excelling in Science and Math.

For many learners as well as educators, reconfigurable equipment such as hobby and learning grade Robots have become a natural medium to learn and experiment. Robots are a special category of machines distinguished by an architecture that allows- multifunctionality, reconfigurability, and re- programmability [1]. The design, building, and operating the robots therefore requires a varied range of skills. Various aspects of robotics therefore become instrument for experiments in logical thinking, programming, and principles of math and physics. To list a few examples:1.2.3.4.Robot navigation and impact of acceleration, deceleration, velocity, and displacement relationships for accurate and time effective motion of robot between two points in a 2D or 3D environment. Correlation between wheelbase, the motor rotation, mobile robot wheel diameter, and total distance covered, steering about the obstacles.

Correlation between wheelbase, the motor rotation, mobile robot wheel diameter, and total distance covered, steering about the obstacles. Levers, gears, kinematic chain, simple machines, and torque to force relationship when lifting objects with the robot arm. Load distribution and the principle of center of gravity when the robot climbs on an incline and decline.178In their research, Ardito et al. and Sana et al. [2,3] demonstrate effective use of robots to teach complex ideas in geometry such as ratios, estimation, and geometry but also finding real life relevance to classroom learning.

and Sana et al. [2,3] demonstrate effective use of robots to teach complex ideas in geometry such as ratios, estimation, and geometry but also finding real life relevance to classroom learning. In their article, Amico, Guastell, and Chella [4] suggest that when using robots for learning, students exhibit better understanding of concepts and higher participation in various topics in physics.Consequently, many educators, parents, and self-motivated individuals participate in various robotics competitive leagues [5,6,7, 8].The application of the robots subdivide them into mobile and non-mobile robots [9,10]. The kinematic control of the robot further classifies it as autonomous or a teleoperated system. However, within competitive robotics most of the challenges are based on mobile robots. At the start of the game, the robot is placed in what is referred to as robot base. During the game, the robot travels to different part of the robot field and completes a variety of tasks.

At the start of the game, the robot is placed in what is referred to as robot base. During the game, the robot travels to different part of the robot field and completes a variety of tasks. The tasks may include pushing, pulling, moving objects, actuating mechanisms, etc. The robot game is usually time bound. The scores assigned are based on total tasks performed within specified challenge duration. In case any exceptions are made, a penalty is assigned. The penalty may negatively impact the total score. In many other competitions, such as robosoccer, sumo-robots, or one-robot events, multiple robots compete as opponents and the scores are determined not only by the robot’s direct accomplishment but also by inhibiting the opponent. Many competitions require autonomous mode, others require teleoperated mode and some have both the rounds.

Many competitions require autonomous mode, others require teleoperated mode and some have both the rounds. World Robotics League follows a progressive approach, and the challenges are based on specific skills.This paper is an outcome of efforts from our Robotics team – ROBOMakers, from North Texas. The team has participated in various competitions since last 4 years and has been recognized for technical excellence. While preparing, the team experimented with many kits, designed and manufactured many parts using the various additive and subtractive freeform fabrication processes. In this paper the team is going to provide a survey of different competitions, kits and their limitations as they are available.

In this paper the team is going to provide a survey of different competitions, kits and their limitations as they are available. The team will provide various case studies as they pertain to limitations with the equipment and usage of Solid freeform fabrication for functional and structural design improvements, adjustment of form factors, load optimizations increasing reliability and customizations for end effectors for the task.2. A brief survey of various competitionsIn this section we are describing popular robotics leagues. Competitive Robotics leagues offer challenges in land, underwater and Aerial Robots. This paper describes application of SFF technology particularly in context of on-land, wheeled robots. Most of the competitions are either completely or in part restrictive about the manufacturers and kits to be used. The format of competition can range from just one set of challenges to structured and systematic progression.

The format of competition can range from just one set of challenges to structured and systematic progression. Table below summarizes a brief overview of various leagues.Table 1 : Brief description of popular competitive robotics leagues179CompetitionWorld Robot OlympiadManufacture Specific Kit Lego with other limited custom partsStructured LearningNoBEST RoboticsVEX kit and Other specified materialsLimitedNo restrictionsYesWorld Robotics LeagueNoFIRST Lego LeagueOnly LegoFIRST Tech Challenge and FIRST Robotics ChallengeREV, Modern, Servocity, Gobilda and some custom partsNoNoLimitedRobocup, JuniorNoNoRobo OneVEX Robotics CompetitionOnly VEXNoNational Underwater RoboticsNoNo3.0 Architecture of a Competitive RobotA typical competitive robot has three main parts:1. The vehicleShort descriptionCompetitions are in four different categories: Regular, Open, Football, and ARC.

The vehicleShort descriptionCompetitions are in four different categories: Regular, Open, Football, and ARC. Specific kit comprising of wood, PVC, DC motors, VEX Hub, wires, motor drivers, and other parts is provided annually. Challenges are offered through the year with emphasis on each participant learning and quantifying progress in structured manner leading to participating in larger team and complex challenge settings. Robotics is just one part of league competitions. It involve designing and programming using only LEGO Mindstorms, NXT, and SPIKE Prime robots to complete tasks based on the year’s theme. Robotics is just one part of league competitions. It involve designing and programming using only league prescribed items to complete tasks based on the year’s theme. The robots compete in one of three main leagues: Soccer, Rescue, or Dance varying in complexity from line-following on a flat surface to negotiating paths through obstacles on uneven terrain.

The robots compete in one of three main leagues: Soccer, Rescue, or Dance varying in complexity from line-following on a flat surface to negotiating paths through obstacles on uneven terrain. Based on bipedal humanoid robots. The competition includes an autonomous stage, followed by one-on-one matches. A challenge is announced annually and the participants design, build, program, and drive in part autonomous and part teleoperated mode. Requires creation of remotely operated underwater vehicle tethered to the surface and manually controlled through underwater camera telemetry. A multitude of end-effectors are used to complete several different tasks180Reference[11][16][15][12][13,14][17][19][18][20]2. The End effector 3. The Controller and power supplyThe vehicle, as name implies, is a mobile platform that takes the robot from one point to another. Additionally, for the objects to be moved and manipulated, end effectors are attached to the robot.

Additionally, for the objects to be moved and manipulated, end effectors are attached to the robot. As described in Figure 1, the end effectors can be grippers, collection tray or launcher. The end effector itself may be connected to an arm. While the vehicle itself may be capable of manipulating the end effector to the desired location and orientation, arm connected to the end effector provides additional articulation.Figure 1 : Architecture of Competition RobotThe wheels as well as the arm and manipulator are connected to DC or Servo motors. Certain competitions require the robots to complete the tasks in autonomous manner [12,15], whereas in other competitions[11,13-20] the robot may be teleoperated, or have two-part events teleoperated, as well as autonomous. A robot controlled in teleoperated mode communicates with remote via WiFi, Bluetooth or similar wireless electronic communication.

A robot controlled in teleoperated mode communicates with remote via WiFi, Bluetooth or similar wireless electronic communication. Motors are powered by an onboard battery and are connected to the controls and communication module. Robots functioning in autonomous mode use multiple sensors such as touch, ultrasonic or vision modules connected to them. Based on the inputs from the sensor, a program in the onboard computer maps the environment and navigates the robot to accomplish the missions.Figure 2, Figure 3 and Figure 4 describe a representative set of challenges offered by different leagues recently. World Robotics League describes a challenge in 3D space. The robot needs to181travel from left to right.

World Robotics League describes a challenge in 3D space. The robot needs to181travel from left to right. As the robot travels, it starts on a platform then in the middle section, it travels on two beams then lands on the platform on the other side.Figure 2 : World Robotics League, Challenge 2 and Challenge 3 of Season 2018[15]Figure 3 : FTC, Challenge Ultimate goal, year 2020-2021[21]182Figure 4 : Vex Challenge Tipping point, year 2021-2022[22]For the FTC challenge, during the Autonomous period, robot needs to deliver Wobble Goals to the randomized Target Zone and then Launch Rings. During the teleoperated mode robot needs to collect and Launch Rings. In the VEX challenge Robot needs to scoring rings, move mobile goals to alliance zones, and elevate on Platforms. It is apparent from the modes of challenges that the missions require quick testing and verification of many different ideas by the teams.4.

It is apparent from the modes of challenges that the missions require quick testing and verification of many different ideas by the teams.4. Description of the kitsThe configurability of the robot kits is required to be able to build and test various mechanisms and subsystems within robot. Popular manufacturers for the kits include Lego[23], goBILDA[24], ServoCity[25], Tetrix[26], REV[27], AndyMark[28], and VEX[29]. The kits include various functional and structural arts including DC motors, servo motors, wheels, metal channels, extrusions, screws, nuts, gears, chains, sprockets, conveyor belt, brackets, connectors, and more.As described in Figure 5 the reconfigurable robot building kits fall into two primary categories, channels with patterned holes, and extrusions. The channels with patterned holes have a set of holes that are arranged in a sequential array. The channels can be attached to each other along the patterned holes.

The channels with patterned holes have a set of holes that are arranged in a sequential array. The channels can be attached to each other along the patterned holes. The patterned hole also allows assembly of the channels along different angle. The angle between the channels is limited to the pitch of the holes. The functional components such as gears, sprockets, and pulleys are connected to shafts via blocks or bearings at the patterned holes. Different elements of the channel-based robots are restricted to linear and the circular pitches of the channel holes.The channel-based kits have a pattern of holes to assemble the structural as well as functional parts. A specific pattern is available to assemble the channels in line or at an angle. Additionally, the adapters for motors, shafts, etc. are designed to adapt to the hole patterns in the channels. The hole design in channels vary based on the manufacturer. The same applies to extrusions.

are designed to adapt to the hole patterns in the channels. The hole design in channels vary based on the manufacturer. The same applies to extrusions. In addition to the structural elements, the hardware between the kits may be incompatible. The DC motors are the strongest motors provided and are usually paired with wheels and/or the mechanisms of the robot. Hex and Servo motors are typically used for mechanisms and are used for driving gears,183chain, sprocket, and a conveyor. The total motors and drives may not be adequate to build a robot, teams usually end up buying additional parts.Figure 5 : Sample of robots built using extrusions and different channelsExtrusion-based kits are similar to most of the industrial extrusion systems with brackets and angles to connect the parts and blocks for mounting motors, shafts and gears and pulleys. Assemblies and subassemblies that are based on extrusion are governed by the size and pitch of the extrusion.

Assemblies and subassemblies that are based on extrusion are governed by the size and pitch of the extrusion. Similar to the channel-based subassemblies, the extrusion-based assemblies are restricted by the pitch of the extrusion system.The interchangeability of different kits is extremely limited. What further limits this ability is the shaft geometry of motors and other rotational components. A team may primarily design and manufacture the robot based on one kit, however, due to the inherent form factor advantage, may want to use the items from a different kit. The team may design and manufacture an adapter that allows use of components from two different kits. Similarly, gears and other such similar elements are available in the matching pattern of the channel or extrusion system. While the robot chassis and arm may be easily built with the extrusion or channel-based elements, end effectors need to be optimized based on the task.

While the robot chassis and arm may be easily built with the extrusion or channel-based elements, end effectors need to be optimized based on the task. The channel and extrusion-based systems become extremely restrictive.In following sections, we will describe different scenarios where 3D printing is used for optimizing and improving overall performance of a competition robot.5. Design and optimization Examples using the 3D printing5.1 Adapters for different Robotic kits184Figure 6 : Custom adapter designed to assemble extrusion and channel based structural elementsFigure 7 : Attaching 3D printed adapter along in line and perpendicular configurationAs described in one of the earlier sections, the ability to interchange parts from different kits is extremely limited. A team may design and manufacture the robot based on one kit but build other subsystems such as drives from another system either due to the form factor advantage or the availability.

A team may design and manufacture the robot based on one kit but build other subsystems such as drives from another system either due to the form factor advantage or the availability. Figure 6 and Figure 7 show a set of parts that are designed and 3D printed to assemble parts from channel and extrusion system into a common assembly. The part is manufactured by Fused Deposition Modeling (FDM), using Orange Acrylic Styrene Acrylonitrile (ASA) printed with T40 tip. Total time required to manufacture the part is 20 minutes.5.2 Optimization of Gripper End effector185Figure 8 : Impact on the Robot due to kit provided vs custom designed and 3D printed parts to build end effectorsFigure 9 : End effector built using 3D printed partsFigure 8 describes a robot with a gripper end effector. The end effector of the robot is built to pick up objects approximately 2inches in diameter. The robot is built using the items provided in the channel-based kit.

The end effector of the robot is built to pick up objects approximately 2inches in diameter. The robot is built using the items provided in the channel-based kit. The system has two griping fingers each actuated by a servo motor. The end effector is mounted on a parallel bar-based arm. When built with the kit provided elements, the size of the end effector becomes extremely large. Usually such an end effector is optimized by using two spur gears driven by a single motor. The engagement of the gear teeth not only allows transfer of torque to both the gears, but also synchronized motion of the two fingers. The gear assembly each requires multiple bearings and mounting blocks, shaft and the gear itself. With the maximum robot weight limitation of 45lbs, the overall weight of the gripper itself is approximately 3.5lbs. Additionally, when the robot arm extends, the center of gravity of the robot shifts beyond the wheelbase, making the robot unstable.

Additionally, when the robot arm extends, the center of gravity of the robot shifts beyond the wheelbase, making the robot unstable. The challenge item to be lifted with the help of the gripper is round. The contacting surface it planar and doesn’t comply with the object to be lifted, thus reducing the reliability of the mission.186As shown in Figure 9 an alternate gripper is manufactured using 3D printing. The total number of parts is significantly reduced. The gripper is driven using a single motor. The gripper is designed to comply with the shape of the object, hence increasing the total contact. The gripper weight is significantly lesser (.37lb). The part is manufactured by FDM, using Orange (ASA) printed with T40 tip. Total time required to manufacture the two parts is 37 minutes.5.3 Cable organization, Shock and vibration reduction using custom suspension for controllera. Empty space under the ramp b.

Total time required to manufacture the two parts is 37 minutes.5.3 Cable organization, Shock and vibration reduction using custom suspension for controllera. Empty space under the ramp b. Hub and battery on the side of the robotFigure 10 : First pass installation of controls and power module.In one of the challenges robot mechanism needs to launch objects. Clear obstruction free surface is needed for smooth launching of objects. As shown in Figure 10. that leaves only sides of the Robot available to install Controls and battery. The hubs installed on the side of the robot are very bulky and take up a lot of space. Additionally, they make the load distribution asymmetric. Mounting the hub under the ramp using screws would restrict the rings from going up smoothly. There is empty space available below the ramp where the battery and controls can be easily accommodated. The problem with this arrangement is: (1) inaccessibility and (2) unreliable wire connection.

There is empty space available below the ramp where the battery and controls can be easily accommodated. The problem with this arrangement is: (1) inaccessibility and (2) unreliable wire connection. A custom cartridge-like housing is designed for a better organization of wires and reduce overall tension on the wires.The control system and the electronics is extremely susceptible to mechanical shocks. The robot turns quickly, travels an uneven terrains, and bumps into objects for a better alignment. The vibration isolation of the control system, receiver and the power supply are required. The kit as such does not provide any fasteners, similar damping or isolating mechanisms3D printed housing is described in Figure 11, instead of on the side, the hubs and battery could go under the ramp with 3D printed part. Pockets are designed for the hub and battery. Features for assembling and guiding the wires are added.

Pockets are designed for the hub and battery. Features for assembling and guiding the wires are added. Silicone based shock absorbers and features to tie the wire for stress release are added to the fixture. Platform with the integrated mechanism isolates the mechanical shocks. The part is manufactured by FDM, using Orange (ASA) printed with T40 tip. Total time required to manufacture the part is 41 minutes.187Figure 11 : Custom designed and 3D printed cartridge base for Robot control and power moduleFigure 12 : Hub and power supply installed on the custom base188Figure 13 : Suspension housed in a telescopic arrangementIn another experiment to improve the ride quality, team built custom Linear Suspension system. As described in Figure 13, the suspension system contains two cylinders that enclose a spring. Both cylindrical enclosures have threaded holes to attach screws to mount to robot.

As described in Figure 13, the suspension system contains two cylinders that enclose a spring. Both cylindrical enclosures have threaded holes to attach screws to mount to robot. However, a basic kinematic analysis suggests that, the suspension system should have additional degree of freedom to move. The thickness of the spring is increased, and the telescopic enclosure is removed.Figure 14: Modified suspension system with additional degrees of freedom.Figure 15: Hub installed via modified suspensionFigure 14 describes the modified suspension system. As shown in Figure 15, the part is manufactured on an Ender 3 Pro using Pla+ material. The total weight of the mechanism is 1 gram. With the telescopic system removed and the thickness of the spring adjusted to match the acting load, the hub ride quality is improved. However, the large mechanical shocks make the part189susceptible to stresses.

However, the large mechanical shocks make the part189susceptible to stresses. The final solution as described in Figure 11 and Figure 12 uses the 3D printed base and an off the shelf silicone suspension.5.4 CameraFigure 16 : The first pass mount for the camera and the camera.The camera is used in the autonomous mode of the robot to analyze the field using machine learning principles. The field of view and distance are critical for proper detection of the features. The camera mount that comes with the kit is about 4 inches in height. Required height should not exceed 2 inches. As shown in Figure 16, in the first revision, the mount for the camera is built using components from the channel system. The limitations of the mounting hole positions still prevent the camera height to exceed 2 inches. The parts from channel system take up a lot of space.Figure 17 : 3D printed custom camera mountA custom mount as shown in Figure 17 is designed and 3D printed for the camera.

The parts from channel system take up a lot of space.Figure 17 : 3D printed custom camera mountA custom mount as shown in Figure 17 is designed and 3D printed for the camera. The 3D printed part is designed precisely to the alignment and with the mounting. This allows better space190utilization along with repeatable and accurate camera performance. The part is manufactured by FDM using Nylon 6. Total time required to manufacture the part is 32minutes.5.5 Custom GearsFigure 19: 3D printed custom bevel gearBevel gears are used for changing the direction of motion and usually arranged in an open configuration (Figure 19). It was quickly realized in our applications that the gears that came with the base kit had small teeth and had limited engagement. This led to frequent slipping between the gears. As shown in figure 19, a bevel gear was designed and then manufactured using an Ender 3 Pro. It was created with PLA+ filament and weighs 4 grams.

This led to frequent slipping between the gears. As shown in figure 19, a bevel gear was designed and then manufactured using an Ender 3 Pro. It was created with PLA+ filament and weighs 4 grams. The 3D printed gear had deeper and wider teeth which reduced the instances of gear slip.5.6 Other opportunities to use 3D printed partsThe ideas discussed in this section were considered during various competitive seasons but not implemented, since other alternative solutions were find. However, they are discussed here due to relevance.5.6.1 Solenoid and kinematic chains191Figure 18 : A custom solenoid activated kinematic chainWith the motor being the primary drive where motion emanates, rack and pinion and kinematic chain mechanisms are used for the conversion of the rotary motion to linear motion. Actuators such as linear solenoids can be used. Pneumatic systems can be used alternatively; however, many leagues do not allow using pneumatic systems.

Actuators such as linear solenoids can be used. Pneumatic systems can be used alternatively; however, many leagues do not allow using pneumatic systems. Custom 3D parts can be designed, and 3D printed to build various kinematic chains. Figure 18 describes a kinematic chain actuated by asolenoid. The solenoid is placed remotely from point of actuation. The point of action stroke is amplified using kinematic chain. To keep the weight low and to maintain the accuracy of the link length, the two links are custom designed and can be fabricated using 3D printing.5.6.2 Custom Pulley for Launching mechanismThe launching mechanism launches a set of disks by spinning two flexible disks in CW and CCW directions respectively. The launch accuracy is dependent upon the speed and the torque of the motor. Space and torque requirements restrict the mechanism to be based on two motors and the speed is synchronized using controllers.

Space and torque requirements restrict the mechanism to be based on two motors and the speed is synchronized using controllers. Unfortunately, most of the motors used are hobby grade gearmotors and with limited speed monitoring using the feedback. The placement of the two motors close to the launch point also makes the mechanism heavy. The weight from the motors can be shifted away from the launch point a timing belt and pulley arrangement. The kit provided pulleys are large and limit the extent of the engagement from the compression wheels. Using 3D printing, a smaller custom pulley can be printed. A mechanism based on gears and pulleys is made to synchronize the two spinners, not only reducing the total part count, but also increasing the accuracy of the system6. Conclusion/SummaryCompetitive robotics offers many learning opportunities and various mechanisms can be built and tested using Robotics kits provided by various manufacturers.

Conclusion/SummaryCompetitive robotics offers many learning opportunities and various mechanisms can be built and tested using Robotics kits provided by various manufacturers. Limitations of such kits can be overcome in cost effective manner by custom design and 3D printing various parts. This paper192demonstrates the same with different examples. The team, in near future will make catalog of parts available to other learners, and participants to use and invite other participants to contribute to the catalog.References:1. Howie M. Choset, Kevin M. Lynch, Seth Hutchinson, George Kantor, Wolfram Burgard, Lydia Kavraki, Sebastian Thrun, Ronald C Arkin Principles of Robot Motion: Theory, Algorithms, and Implementation, , MIT Press, 2005 - Technology & Engineering, ISBN- 13: 978-02620332752. Ardito, G.; Mosley, P.; Scollins, L. We, robot: Using robotics to promote collaborative and mathematics learning in a middle school classroom. Middle Grades Res. J. 2014, 9, 73–88.3.

Improving Children’s Coordinative Skills and Executive Functions: The Effects of a Football Exercise Program. Percept. Mot. Skills 2016, 122, 27–46, doi:10.1177/0031512515627527.7. Bredenfeld, A.; Hofmann, A.; Steinbauer, G. Robotics in Education Initiatives in Europe—Status, Shortcomings and Open Questions. In Proceedings of the SIMPAR 2010 Workshops International Conference on Simulation, Modeling and Pro-gramming for Autonomous Robots, Darmstadt, Germany, 15–16 November 2010; pp. 568–574.8. Dwivedi, R; Kumar A; Babu, B; Grandhi, N; Meka,R; Ahuja, V;, The Role of Competitive Robotics in Providing Context to Classroom Learning and Technical Skill Development in School Age Students—A Survey of Current Avenues, Assessment, and Path Forward with Systematic Implementation MPDI Journal of Education Sciences, Volume 11, Issue 4, April 2021.9. Dwivedi, R.; Dwivedi, S.

viewed PaperThe Impact of the Printed Part Geometry on the Shrinkage and the Density of 316L Stainless Steel Parts Printed by FFF/FDM TechnologySuleiman Obeidat*, Joe Nervis Jr., Junkun MaDepartment of Engineering Technology, Sam Houston State UniversityCorresponding author: obeidat@shsu.eduAbstractAdditive manufacturing has been used extensively for the last decade in making different parts of different complexities. One of the additive manufacturing methods discussed in this work is Fused Filament Fabrication/ Fused Deposition Modeling (FFF/FDM). FDM is used to fabricate stainless steel using BASF Ultrafuse 316L Metal 3D printing filament which is made of 80 w. % 316L stainless steel particles in polymer base. We study the impact of the part geometry on the part shrinkage, apparent and relative density of 3D printed parts (green and sintered parts). Four basic geometries are printed: rectangular blocks, cylinders, spheres, and cones.

Four basic geometries are printed: rectangular blocks, cylinders, spheres, and cones. These geometries have the same volume with the same height in the printing direction for the cylinders, rectangular blocks, and the cones. It is found that the relative density (apparent density/theoretical density) of approximately 99.0 % is achieved in the rectangular block compared to approximately 96.65 % for the cylinder, 98.0 % for the cone, and 95.8 % for the sphere. The highest shrinkage of 47.9 % takes place in the cylinder while the shrinkage in the rectangular block is approximately 44.42 % as a maximum.Keywords: Additive manufacturing; FFF/ FDM; 316L stainless steel.IntroductionAdditive manufacturing is one of the operations used to make parts layer by layer instead of cutting materials to shape the part made as in the machining processes.

In the last two decades, the industry has been using the additive manufacturing extensively in making complicated parts that can be used in different applications such as biomedical implants, heat engines and other applications that need complicated and strong parts with intricate details. But as any other manufacturing processes, making fully defect free parts is not possible. These defects such as pores and cracks in the printed part impact the mechanical properties of the part. These pores and cracks come from the way the part is manufactured, the manufacturing process parameters and the complexity of the part which include its shape. Different researchers studied different factors that affect the mechanical properties of the parts printed. [1] Liu et al. (2020) and [2] Gong et al. (2019) compared between the parts made by FDM and those ones made by Selective Laser Melting process.

[1] Liu et al. (2020) and [2] Gong et al. (2019) compared between the parts made by FDM and those ones made by Selective Laser Melting process. [1] stated that FDM technology followed by de-binding and sintering is a more energy- saving process to fabricate metal parts compared to Selective Laser Melting process. They used the FDM technology to produce 316L stainless steel and they concluded that the parts made using FDM are weaker than the parts made using Selective Laser Melting but at the same time they stated that denser products can be made with different printing parameters and improved sintering process.[3] Schumacher and Moritzer (2021) used FDM technology and the BASF 3D Printing1870Solutions GmbH material Ultrafuse 316L to make metallic parts. They studied the effect of material specific FDM processing parameters on the properties of the sintered parts (white parts).

They studied the effect of material specific FDM processing parameters on the properties of the sintered parts (white parts). They used the density of the green parts produced as an evaluation criterion to assess the white part quality. [4] Abe et al. (2021) studied the effect of layer directions on internal structures and tensile properties of 17-4PH stainless steel parts fabricated by FDM. They found that the as sintered specimens printed with its direction perpendicular to the tensile direction achieved the highest ultimate strength. Also, they found that the linear shrinkage is higher in the printing direction than in the other directions. [5] Wang et al. (2019) studied the effect of raster orientation and extrusion width in the FDM process on the porosity and mechanical behavior of 3D printed parts. They used X-ray computed tomography (XCT) to characterize the pores inside the printed structure to predict the mechanical properties of the parts printed. [6] Quarto et al.

They used X-ray computed tomography (XCT) to characterize the pores inside the printed structure to predict the mechanical properties of the parts printed. [6] Quarto et al. (2021) used Analysis of Variance to study the effect of different printing parameters on the density, porosity and shrinkage of the 316L stainless steel parts printed using fused deposition modeling (FDM). [7] Patti et al. (2021) studied the thermo- mechanical properties and the effect of porosity on those properties of 316L stainless steel printed using FDM at different infill densities. [8] Hassan et al. (2021) investigated the microstructure and the mechanical properties of the 316L stainless steel at different values of part orientation, extrusion velocity and layer height. They found that there is a minor effect of the extrusion velocity and the layer thickness on the porosity and the grain size. [9] Tosto at al. (2021) studied the properties of Ultrafuse 316L BASF and 14-4 PH Makerforged printed using FDM.

[9] Tosto at al. (2021) studied the properties of Ultrafuse 316L BASF and 14-4 PH Makerforged printed using FDM. They investigated the mechanical properties of the printed parts and they related them with the raster bonding and the voids present in the green and sintered parts. [10] Rosnitschek et al. (2021) studied the effect of the part geometry and the infill degrees on the mechanical properties of 316L stainless steel fabricated by FDM. [11] Maconachie at al. (2020) confirmed that the mechanical behavior of the gyroids fabricated by fused deposition modeling is affected by the geometry and the topology of the gyroids rather than the manufacturing effects. They used in their study acrylonitrile butadiene styrene (ABS) to print those parts. [12] Moritzer et al. (2021) studied the effect of the strand geometry and nozzle size on the density of the green parts and the mechanical properties of the sintered 3D printed parts.

[12] Moritzer et al. (2021) studied the effect of the strand geometry and nozzle size on the density of the green parts and the mechanical properties of the sintered 3D printed parts. Our objective is to study the effect of the geometry of the part fabricated using Fused Filament Fabrication/ Fused Deposition Modeling (FFF/ FDM) on the apparent and relative density and the shrinkage percentage of the sintered parts. We used the BASF Ultrafuse 316L Metal 3D printing filament which is made of 80 w. % 316L stainless steel particles in polymer base. It is up to the authors knowledge that this topic has not been studied in this way. General geometries such as rectangular bocks, cylinders, cones, and spheres are printed using (FFF/ FDM). BASF Ultrafuse 316L Metal 3D printing filaments are used to make these parts.Materials and Methods: Parts FabricationThe parts are made using (FFF/FDM) process using MakerBot Method X 3D printer.

BASF Ultrafuse 316L Metal 3D printing filaments are used to make these parts.Materials and Methods: Parts FabricationThe parts are made using (FFF/FDM) process using MakerBot Method X 3D printer. The process starts with creating a CAD model using a computer aided design (CAD) software then convert this file into (.stl) format which is compatible with these types of printers. The filament material we use is 316L Stainless steel powder of 80 w. % mixed with 20 w. % resin. Fig. 1 shows this printer. Filaments of 1.75 mm diameter have been used. The four basic shapes used are rectangular blocks, cylinders, spheres, and cones. The building plate is prepared by scraping any1871residues from previous prints and then adding Magigoo Pro Metal 3D Bed Adhesion Solution for BASF Ultrafuse 316L. As shown in Fig. 1, this printer has two extruders; one of them is for building material which is 316L stainless steel, and the other is for support material.

As shown in Fig. 1, this printer has two extruders; one of them is for building material which is 316L stainless steel, and the other is for support material. Here we did not use support material since we don’t have over hanged parts and there is no possibility of bending the parts printed. We used one extruder of LABS extruder type which is used mainly for metals. The printing parameters used are all the same in all the prints to be consistent and to only investigate the effect of the part geometry on the shrinkage and the density of the part printed.5 cmFig. 1: Makerbot Method X 3D printer.We printed 3 cylinders, 3 rectangular blocks, 3 spheres, and 3 cones. The parts built are created in a way that their volumes are the same in all the prints. A cylinder of 15 mm diameter and 20 mm height has been chosen to start with. The volume of this cylinder is 3534.3 mm3.

A cylinder of 15 mm diameter and 20 mm height has been chosen to start with. The volume of this cylinder is 3534.3 mm3. We chose the height to be the same for the cylinders, the cones, and the rectangular blocks and then we calculated the other dimensions based on a constant volume (3534.3 mm3) and a height of 20 mm. The rectangular block is chosen to be of a square base. Table 1 shows the dimensions of the parts to be printed. After the fabrication process, the parts are de-bound and sintered to improve their mechanical properties. Some of the printing parameters used are layer thickness: 0.2 mm, extrusion temp.: 245 ℃, chamber temp.: 85 ℃, infill percentage: 100 %, infill pattern: linear, and the travel speed: 250 mm/s.

Some of the printing parameters used are layer thickness: 0.2 mm, extrusion temp.: 245 ℃, chamber temp.: 85 ℃, infill percentage: 100 %, infill pattern: linear, and the travel speed: 250 mm/s. All the parameters are the same in all samples.Table 1: Dimensions of the parts printed.PartRectangular block Cylinder Cone SphereVolume (mm3) 3534.3 3534.3 3534.3 3534.3Height (mm) 20 20 20 NADiameter (mm) NA 15 25.98 18.9Width (mm) 13.29 NA NA NASurface area (mm2) 1416.448 1295.25 1345.616 1121.6391872Materials and Methods: De- Binding and Sintering ProcessThe samples were sent to DSH technologies, LLC for the thermal de-binding and sinteringprocess. The green parts and the sintered ones are shown in Fig. 2.15 mm15 mmFig. 2: 316L stainless steel printed green parts (left) and sintered parts (right).The dimensions and the densities have been measured for the green parts and the sintered parts.

2.15 mm15 mmFig. 2: 316L stainless steel printed green parts (left) and sintered parts (right).The dimensions and the densities have been measured for the green parts and the sintered parts. Three samples have been printed and sintered of each geometry shown in Fig. 2. The dimensions have been measured using a regular caliper and the density has been measured using Archimedes method. Torbal AGCN220 Internal Calibration Analytical Balance, 220 g x 0.0001 g balance has been used in the process in addition to Torbal Density Kit for AGCN Scales.Results and DiscussionAs mentioned above the density has been measured using the Archimedes method. Apparent density has been calculated using equation 1 [13]. Fig. 3 shows the apparent density of the green parts.

Apparent density has been calculated using equation 1 [13]. Fig. 3 shows the apparent density of the green parts. It shows that the rectangular block apparent density is higher than that of the cylinder which in turn is higher than that of the cone and the sphere.(1)𝜌𝜌𝑎𝑎 = 𝜌𝜌𝑤𝑤𝑤𝑤: Apparent density of the part.𝑚𝑚𝑑𝑑 𝑚𝑚𝑑𝑑−𝑚𝑚𝑤𝑤𝜌𝜌𝑎𝑎: Water density at the measurement temperature𝜌𝜌𝑤𝑤𝑤𝑤: dry mass of the part𝑚𝑚𝑑𝑑: wet mass of the partFig. 4 shows the apparent density of the sintered parts. It is noticed that the sintered parts have 𝑚𝑚𝑤𝑤 almost the same trend as the green parts for some samples where the highest apparent density is for the rectangular block, then cylinder, sphere, and cone. In other samples, the cone or the cylinder have the highest apparent density. The apparent density of the rectangular block is the highest between them all. This might be because the rectangular blocks have the largest surface area as shown in table 1.

The apparent density of the rectangular block is the highest between them all. This might be because the rectangular blocks have the largest surface area as shown in table 1. Having a high surface area means that the rate of heat transfer is higher which makes the de-binding and the sintering process more efficient which in turn improves the part density. The surface areas of the cylinder and the cone are close to each other which explains why the apparent density of those two parts are close to each other. The surface area of the sphere is the smallest, but its apparent density is close to that of the cylinder in some samples. The bars shown in Fig. 4, Fig. 5, and Fig. 6 are the standard deviation of the measured values. From these1873bars, it is concluded that more research is needed with different shapes, bigger size parts, and different printing parameters.Fig. 3: Green parts apparent densityFig. 4: Sintered parts apparent densityFig.

3: Green parts apparent densityFig. 4: Sintered parts apparent densityFig. 5 shows the relative density (apparent density/theoretical density) of the 316L stainless steel. The theoretical density used here is 7.8 g/cm3. Fig. 5 shows that a range of about 95.0 % to 99.0 % relative density in the rectangular block is achieved compared to a maximum of 96.65 % for the cylinder, about 98.0 % for the cone, and about 95.8 % for the sphere . This indicates again that a more relative density is achieved with higher surface area of the part printed but more investigation is needed with some other shapes and bigger volume parts. The shrinkage percentage has been measured, too. The volume of the sintered parts is compared with the volume of the green1874parts. Fig. 6 shows the shrinkage percentages for the different shapes under study.

The volume of the sintered parts is compared with the volume of the green1874parts. Fig. 6 shows the shrinkage percentages for the different shapes under study. It is noticed that the highest shrinkage takes place in the cylinder with 47.9 %, then the sphere with a maximum shrinkage of 46.63%, then cone with 45.77 %, and finally the rectangular block with a maximum of 44.42 %. This also is justified by the highest surface area of the rectangular block compared to the other geometries.Fig. 5: Relative density of the 316L SS after sinteringFig. 6: Volume shrinkage of the sintered partsConclusionsFFF/ FDM technology is used to study the impact of the part geometry on the part shrinkage and the apparent density. 316L stainless steel parts at specific printing parameters are built in different shapes such as rectangular blocks, cylinders, cones, and spheres. The parts built are of the same volume but of different surface areas.

The parts built are of the same volume but of different surface areas. We sintered the samples to obtain the metallic parts1875of high density. Archimedes method is used to measure the apparent density of the green and sintered parts. Approximately 99.0 % relative density is achieved as a maximum in rectangular blocks while a 93.42 % is achieved in the cones as a minimum. The highest shrinkage of 47.9 % takes place in the cylinder while the shrinkage in the rectangular block is about 44.42 % as a maximum. One of the reasons for this result is the highest surface area of the rectangular blocks compared to the other geometries. Research is underway to investigate the impact of the raster angle, the overlap between the rasters, and to investigate different shapes and sizes on the relative density.

Research is underway to investigate the impact of the raster angle, the overlap between the rasters, and to investigate different shapes and sizes on the relative density. It is expected that the overlap between the rasters would improve the relative density.REFERENCES[1] Liu, Bin; Wang, Yuxiang; Lin, Ziwei; Zhang, Tao, “Creating metal parts by Fused Deposition Modeling and Sintering”, Materials Letters, Volume 263, 15 March 2020, 127252.[2] Haijun, Gong; Snelling, Dean; Kamran, Kardel; and Carran, Andres, Comparison of Stainless Steel 316L Parts Made by FDM-and SLM-Based Additive Manufacturing Processes, JOM, Vol. 71, No. 3, 2019, https://doi.org/10.1007/sl1837-018-3207-3 © 2018 The Minerals, Metals & Materials Society.[3] Schumacher, Christian; and Moritzer, Elmer, “Steel Parts Produced by Fused Deposition Modeling and a Sintering Process Compared to Components Manufactured in Selective Laser Melting”, Macromolecular Symposia, Vol.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperImpact of Vapor Polishing on Surface Roughness and Mechanical Properties for 3D Printed ABSClayton Neff, Matthew Trapuzzano, Nathan B. CraneDepartment of Mechanical Engineering, University of South Florida, Tampa, FLAbstractAdditive manufacturing (AM) is useful when creating complex geometric models and prototypes. However, a well-known drawback is the fact that parts produced by AM methods typically have lower strength and higher surface roughness than traditionally-formed parts. To compensate for this, the surface finish is commonly improved by mechanical finishing or some type of coating. Another widely used surface treatment for ABS components is vapor polishing.

To compensate for this, the surface finish is commonly improved by mechanical finishing or some type of coating. Another widely used surface treatment for ABS components is vapor polishing. In this process, the part is exposed to a solvent vapor that partially dissolves a surface layer and enables smoothing through surface tension-driven flow; it is known to decrease the surface roughness. However, little work has been reported quantifying the surface roughness change or on the mechanical impacts of this processing method. This work compares the strength, ductility and surface finish of vapor-polished ABS tensile specimens of varying thicknesses (1 mm, 2mm, and 4 mm). Results show that elongation at break is improved, while the modulus of elasticity is reduced in thin specimens. The tensile strength is largely unchanged.

Results show that elongation at break is improved, while the modulus of elasticity is reduced in thin specimens. The tensile strength is largely unchanged. The power spectral density for roughness features larger than 20 µm were reduced 10X.IntroductionAdditive manufacturing designates the formal expression for the manufacturing process popularly known as “3D” printing or rapid prototyping. AM processes are best-suited for low volume production runs because no part-specific tooling is required. Many AM processes are based on conventional printing processes such as inkjet and laser printing with stacked layers. The term “3D printing” emphasizes this connection and is commonly applied to the low cost AM processes. Another distinct advantage of AM over traditional manufacturing includes the ability to create highly complex parts of near arbitrary geometry.

Another distinct advantage of AM over traditional manufacturing includes the ability to create highly complex parts of near arbitrary geometry. [1, 2] This advantage essentially allows designers to bypass the “design for manufacturing” step that is crucial to traditional manufacturing methods like casting.One of the most common AM processes resides in material extrusion technologies. This process works most often with thermoplastic material in either a filament or pellet form. A pinch- roller or screw system feeds the material into a heated extrusion head [3]. The thermoplastic material begins to liquify in the heated extrusion head and then is deposited onto a heated platform through the nozzle. The material flow through the nozzle is controlled by the volume of material fed into the nozzle.

The material flow through the nozzle is controlled by the volume of material fed into the nozzle. The pressure drop between the liquifying chamber in the nozzle and the surrounding atmosphere can produce a time delay between changes in the rate of volume input and flow through the nozzle [4]. A computer generates a numerical code for the coordinates of a “sliced” digital file of a prototype or part and these coordinates become the pathways or “roads” for material deposition of part geometry. Once the extruder head deposits the first layer of material,2295the heated platform will descend or the extrusion head will rise one layer height and proceed to deposit the material in a layer by layer fashion. This process again acknowledges the “additive” approach of AM as the extrusion head adds material layer by layer.

This process again acknowledges the “additive” approach of AM as the extrusion head adds material layer by layer. Extrusion technologies depositing layers in this fashion are often termed Fused Deposition Modeling or “FDM” machines, first produced and developed by Stratasys [5].The FDM process is the quintessential AM process because it is the process most people think of first when someone mentions “3D printing.” However, unlike plastic parts manufactured by traditional methods such as casting or injection molding, parts manufactured by the FDM process suffer certain drawbacks. These drawbacks have stifled their use in many of the applications that would benefit from the free-form fabrication that additive manufacturing has to offer. FDM commonly prints with the material of Acrylonitrile Butadiene Styrene (ABS). These parts are typically rough due to the discrete layers and roads of the deposition.

FDM commonly prints with the material of Acrylonitrile Butadiene Styrene (ABS). These parts are typically rough due to the discrete layers and roads of the deposition. They also warp often and have anisotropic mechanical properties significantly worse than their bulk-manufactured counterparts. [6-10] It is these drawbacks that limit applications of the FDM process. Once these limitations are addressed, the FDM process will find wider range of application.One approach to improving the surface finish is through post-processing methods. Many parts are mechanically polished or coated to improve the surface. However, another post- processing technique gaining particular interest for ABS is vapor polishing. Vapor polishing acts as a surface treatment to reduce the surface roughness with the potential to maintain dimensional accuracy without significantly eroding part geometry. [11-13] It is also relatively low-labor process compared to mechanical polishing and coating.

[11-13] It is also relatively low-labor process compared to mechanical polishing and coating. Vapor polishing introduces a solvent solution with a significant vapor pressure (acetone) that absorbs into to the ABS parts and penetrates the polymer. The solvent dissolves the polymer, and the surface tension forces drive a reduction in the surface roughness in a process similar to viscous sintering. This produces a more even, smooth, and better quality surface finish. In prior work on thermoplastic parts, controlled solvent exposure has shown to improve strength of weld-lines when joining ABS geometries as well [14]. The following experiments show the benefits of using vapor polishing as a post- processing method to improve the mechanical properties, as well as the surface finish, of ABS FDM parts.Experimental MethodsFor mechanical testing purposes, parts were printed according to ASTM Standard D638– 10 [15]. The dimensions of the tensile barbell specimens are shown in Figure 1.

The dimensions of the tensile barbell specimens are shown in Figure 1. These barbells were printed on a Stratasys uPrint SE machine in ABS Plus material with various thicknesses of 1, 2, and 4 mm. The parts were printed with the long axis oriented in the z-direction (normal to the print bed) and the ZXY plane – as designated by ASTM F2921 [16]. The z-direction typically has the highest surface roughness and weakest bond strength between deposited layers of FDM fabricated components; therefore, the z-direction was selected as the print orientation. This was done in hopes to reveal the effect of surface roughness and mechanical properties between unpolished and polished samples. Figure 2 below illustrates a sample tensile specimen being printed with included dashed lines as deposition lines where layers bond together. Dissolvable support material was used to accommodate the long slender parts being printed in this fashion.

Dissolvable support material was used to accommodate the long slender parts being printed in this fashion. The parts were printed solid with 100% infill and layer thicknesses of 254 microns. The unpolished2296batch of tensile specimens consist of five parts for each thickness, for a total of 15 parts. The support material was dissolved in caustic soda, then after the conditioning periods described below, pre-experiment dimensional measurements of each part were taken using a digital caliper after environmental conditioning explained below.A Veeco Dektak 150 profilometer was utilized in order to reveal the surface topology of the tensile specimens. A typical contact profilometer uses a stylus to contact a specimen and then move laterally across the surface to register the profile changes in the vertical direction. The Veeco Dektak 150 used has a stylus with a radius of 12.5 μm and a resolution of 278 nm.

The Veeco Dektak 150 used has a stylus with a radius of 12.5 μm and a resolution of 278 nm. For this study, profilometry scans were oriented along the z-direction of the parts with a scan length of 5 mm in 60 seconds with 10 mg of contact force. All samples were measured to find the average RMS roughness of the “as printed” unpolished parts.Figure 1: Dimensions of tensile specimensFigure 2: Build orientation of a tensile specimen showing the dashed deposition lines and surrounding support materialThree to five samples of each thickness were printed and then set aside to be vapor polished. The vapor polishing was done on three samples (one of each thickness) at a time. Two napkins of the same size were soaked with 5 ml of acetone and placed around the perimeter inside a 1 L (32 oz.) high-density polyethylene container. The tensile specimens were hung from a metal fixture inside the container and the container was sealed for 45 minutes.

The tensile specimens were hung from a metal fixture inside the container and the container was sealed for 45 minutes. Figure 3 shows a schematic of the vapor polishing method. The polished parts were allowed to completely dry for 5 days (120 hours) under ASTM D618 procedure A conditioning requirements. Five days was allotted for drying to ensure the acetone evaporated out of the polished tensile specimens before testing. Figure 4 shows the mass of a sample 4 mm tensile specimen over time. The average (with standard deviation) weight gain is likely due to residual acetone or water vapor for the different thicknesses: 4 mm specimens have an average of 1.06% (St. Dev. of 0.003) weight gain [g], 2 mm specimens have an average of 2.83% (St. Dev. of 0.004) weight gain [g], and the 1 mm specimens have an average of 5.43% (St. Dev. of 0.005) weight gain [g]. Absolute weight gain is similar for all part thicknesses which is consistent with a surface-mediated process.

Dev. of 0.005) weight gain [g]. Absolute weight gain is similar for all part thicknesses which is consistent with a surface-mediated process. Profilometry and dimensional measurements using a digital caliper were performed again on the post-processed barbells for later comparison of dimensional changes as well as changes in surface finish.Finally, the specimens were mechanically tested in an MTS 858 hydraulic table-top tensile testing unit after conditioning the tensile specimens in accordance to ASTM D 618 Procedure A:229740/23/50 for specimens ≤ 7 mm (0.25”). The procedure calls for the standard laboratory atmosphere of 23°C 10% relative humidity to condition the test specimens for a minimum of 40 hours [17]. The 40 hours minimum was applied for unpolished samples but at least 120 hours was applied to the polished samples to allow for evaporation of the acetone as discussed above. The testing was done in accordance to the ASTM Standard D638–10 [15].

To assess the dimensional impact, dimensional data of the tensile specimens were taken before and after vapor polishing was introduced. Each measurement was repeated at least 3 times at different locations for thickness, width, and length. Table 1 below shows the measured dimensional changes recorded for pre and post-polishing of the samples meanwhile Table 2 depicts the percentage change for the dimensional changes. Table 1 and Table 2 indicate the change of2298thickness is negligible with less than 1% change for the different thicknesses. The length change of the specimens is negligible, meaning that the 45 minute polishing duration is short enough to prevent slumping of the geometry due to the gravitational forces pulling material to the bottom end of the specimen. The only significant change in dimensions appears in the percentage change for the width of the 1 mm specimen.

The only significant change in dimensions appears in the percentage change for the width of the 1 mm specimen. This suggests there may be a small surface effect resultant from smoothing of the bulging layer extrusions. The surface erosion acts more on the 1 mm samples as it effects a larger percentage of the width. Since most of the dimensional changes are less than 1%, the changes remain within the tolerance threshold of the uPrint machine.Surface RoughnessOne of the key parameters of this study evaluates the surface roughness of vapor polishing ABS FDM components. Figure 5 shows an FFT analysis of the surface roughness along the z- direction; this figure shows one sample before and post-processing. In each case, there is a peak corresponding to ~ 3.75 , or ~ 254 μm (the characteristic layer thickness).

In each case, there is a peak corresponding to ~ 3.75 , or ~ 254 μm (the characteristic layer thickness). The difference between the unpolished and polished samples show there was a 10X reduction in the frequency band corresponding to the layer thickness noise and a significant reduction at all dimensions that were apparent in the unpolished sample. Figure 6 below shows SEM photographs of unpolished vs. polished tensile specimens for the various thicknesses. The SEM photos show the substantial impact vapor polishing has on the surface roughness. The unpolished samples have peaks and valleys at each new deposition layer (254 μm layer). However, the polished samples show a large reduction in the artifact. Table 3 shows that polishing sharply decreased RMA surface roughness and variation in the surface roughness.−1mmTable 3: Average roughness changes of post-processed specimensAs-printed (μm)Vapor polished (μm)Ra RqSt. Dev. St.

Dev. St. Dev.37.18 44.4114.55 15.0610.13 12.398.23 8.86± ±± ±± ±y t i s n e D101Peak of layer thicknessUnpolishedPolishedl a r t c e p S r e w o P g o L0.10.010.0010.0001050 150 100 Spatial Frequency (mm-1)200Figure 5: FFT analysis of a sample specimen for the surface roughness along the build orientation (z-direction)2299(a)(c)(e)Figure 6: SEM of: 1 mm unpolished (a), 1 mm polished (b), 2 mm unpolished (c), 2 mm polished (d), 4 mm unpolished (e), 4 mm polished (f)2300(b)(d)(f)Mechanical PropertiesFigure 7 shows the strain to failure, elastic modulus, and ultimate tensile strength vs. thickness with a sample stress vs. strain chart. No significant strength differences were observed between the polished and unpolished samples, however the strain to failure of the polished samples was higher than the unpolished (the difference falls within one standard deviation for all parts except the 2 mm thick parts).

The strain to failure for the 1 mm polished specimens have the largest increase, but the elastic modulus has the largest decrease. Polishing has reduced the elastic modulus in all specimens, but the effect decreases with increasing thickness. The sample stress vs. strain chart supports this comparison as well. The 1 mm polished specimen also has more elongation and even a little yielding before fracture, but a lower value in slope (elastic modulus).1.82.1)%(e r ul i a F o t ni a r t S1.61.41.21.00.80.6UnpolishedPolished) a P G(s u u d o Mlc i t s a l E2.01.91.81.71.61.51.41.3UnpolishedPolished0.41.202460246Thickness (mm)Thickness (mm)) a P M( h t g n e r t S e l i s n e T.

t l222018161412UnpolishedPolished) a P M(s s e r t S2520151051 mm Unpolished1 mm PolishedU1002460 0.0000.0050.0100.015Thickness (mm)Strain (mm/mm)Figure 7: Mechanical property charts for unpolished vs. polished samples: strain to failure vs. thickness (top left), elastic modulus vs. thickness (top right), ult. Tensile strength (bottom left), stress vs. strain curves (bottom right)DiscussionAcetone vapor polishing has a noticeable effect on the dimensional characteristics of ABS components of 1 mm thickness but negligible as the thickness increases. However, the noticeable effect could be minimized if the duration for polishing is reduced for thinner parts or less solvent is used. The outer texture of FDM parts have small rounded ridges that bulge out on each layer2301from the cylindrical extrusion being flattened.

The outer texture of FDM parts have small rounded ridges that bulge out on each layer2301from the cylindrical extrusion being flattened. Acetone vapor polishing partially dissolves some of this material on the surface of the part and allows the “high” material to flow into the “low” areas between each layer under surface tension effects. Gravitational and surface energy effects can both contribute to the material flows, especially as the viscosity decreases with increased acetone exposure. In general, if the acetone vapor polishing was done completely, one would think the final location of the surface would be somewhere close to the middle of the high and low points. However, the resulting dimensional change is insignificant for most applications. This is a desired result of any post-processing method.Acetone vapor polishing has a profound effect on the surface finish of an ABS part.

This is a desired result of any post-processing method.Acetone vapor polishing has a profound effect on the surface finish of an ABS part. The average roughness and RMS roughness were both decreased significantly due to acetone vapor polishing. The sole purpose of this effect allows ABS FDM parts to have a smooth surface finish and be more aesthetically appealing to the eye instead of a rough granular one. For surface finish and dimensional tolerances, careful consideration needs to be taken as to the duration of polishing; over-polishing will distort the surface and cause tolerance issues. The process outcomes depend on the total acetone available, surface area of the parts, exposure time, and container volume. Further work is needed to characterize the relationship between the process parameters and the surface roughness outcomes.Vapor polishing also affects the mechanical properties of ABS FDM parts.

Further work is needed to characterize the relationship between the process parameters and the surface roughness outcomes.Vapor polishing also affects the mechanical properties of ABS FDM parts. In general, the polished specimen strength is comparable (though slightly higher) than the unpolished specimens, and the elongation to break is increased. This is a desired effect that is hard to obtain with most traditional materials. Thinner components see a larger impact (ductility increased and reduced modulus of elasticity) than their thicker counterparts. This result is most likely due to the relatively larger surface area/volume ratio of the thinner specimens. The property change could be due to the shape change, but we believe it is likely related to the residual weight gain of the treated parts. The additional weight may be due to absorption of a plasticizing agent.

The additional weight may be due to absorption of a plasticizing agent. A plasticizer by definition defines a low molecular weight polymer additive that enhances flexibility and workability but reduces stiffness. [18] In terms of this study the weight gain – either acetone, water vapor, or both – may be acting as plasticizer molecules to occupy interstitial positions between the larger polymer molecules. This would effectively increase the inter-chain distance between the larger polymer molecules. An increase in inter-chain distance would allow further elongation – thus strain before failure – but also comes with the penalty of reducing the intermolecular bonding, thus resulting in lower elastic modulus.An interesting result to note appears in the trends of the ultimate tensile strength of polished versus unpolished specimens.

As mentioned earlier, the 1 mm specimens achieve the highest strength value for unpolished samples but in opposition the 4 mm specimens achieve the highest strength value for polished samples. A possible solution to explain this trend may be the print time between each layer as the tensile specimens were printed in batches of same thickness. Therefore, the time between layers for thicker parts will be longer since the extrusion head has a longer deposition path to follow before it finishes a layer on each specimen. This translates to each layer having longer time to cool resulting in a decrease of adhesion between each layer. On the contrary, thinner specimens would have less time to cool between each layer resulting in greater bond strength. Since the adhesion of each new layer would integrate more into the previous layer, the thinner specimens would be able to reach higher stress levels.

Since the adhesion of each new layer would integrate more into the previous layer, the thinner specimens would be able to reach higher stress levels. Vapor polishing erases this trend as2302the adhesion between each layer during printing becomes less of a factor by intertwining the polymer chains of new layer together; resulting in higher achievable stress levels for thicker specimens.The orientation of the z-direction was chosen for the printed tensile specimens as this is typically the weakest direction and thus the one most likely to benefit from post-processing. Because of the orientation of the printed parts during testing, they essentially have small cracks between each layer oriented normal to the force. This allows brittle fracture to occur as a result of mode I crack opening failure.

This allows brittle fracture to occur as a result of mode I crack opening failure. Acetone vapor polishing fills the cracks on the outside – and depending on the thickness, the whole part – with dissolved material, either shortening the crack length or eliminating the cracks completely. This should result in better mechanically-performing parts, however this effect was insignificant in these tests. It seems unlikely that there would be an impact in other printing orientations.This work considered only the case for a full infill. Parts with partial infill may behave very differently as there are many more interfaces to bond. This may create new opportunities for property enhancement through vapor polishing.

Parts with partial infill may behave very differently as there are many more interfaces to bond. This may create new opportunities for property enhancement through vapor polishing. Partial infill may also introduce new failure mechanisms as thinner surface sections may collapse with excessive acetone.In general, more data needs to be collected, and variations of the post-processing methods (especially those that vary the strength) need to be further explored to determine what the ideal post-processing parameters should be in order to obtain better looking parts that have bulk-like mechanical properties, but still maintain dimensional tolerance.ConclusionsAM-printed FDM components have obvious drawbacks of surface roughness and anisotropic properties. This study investigates the effect of post-processing ABS tensile specimens by subjecting them to vapor polishing in an effort to improve on the drawbacks.

This study investigates the effect of post-processing ABS tensile specimens by subjecting them to vapor polishing in an effort to improve on the drawbacks. Vapor polishing decreases surface roughness vastly while also having an impact on mechanical performance. It was found that vapor polishing has a larger impact on thinner components by increasing strain to failure and strength but decreasing the elastic modulus. Thicknesses above 2 mm show a modest improvement in ductility and strength with a modest decrease in elastic modulus. This study helps to improve the quality of ABS FDM components and with future work can ultimately advance FDM technology.2303References1.2. 3.4. 5.6.7.8.9.10.11.12.I. Gibson, D.W.R., B. Stucker, Additive ManufacturingTechnologies: Rapid Prototyping to Direct Digital Manufacturing. 2010, New York: Springer. Gebhardt, A., Understanding Additive Manufacturing. 2012: Carl Hanser Verlag. Brian N. Turner , R.S., Scott A.

viewed PaperImpact of Zirconia Slurry Doping on 316L Stainless Steel prepared by Laser Powder Bed Fusion for biological/high corrosion applicationsMcKay Sperry*, Jakob Bates*, Taylor Davis*, Tracy W. Nelson*, Nathan B. Crane*Brigham Young University, Provo, UT, USAAbstract Laser Powder Bed Fusion (LPBF) is a versatile additive manufacturing (AM) method used primarily for processing plastics and metals. Stainless steel (type 316L) is a biocompatible metal used extensively for LPBF and commonly for medical applications. Zirconium Dioxide (zirconia) is a common engineering material used in applications ranging from dental implants to oxide dispersion strengthening of metals. This paper presents the process development and results of in- situ deposition of small quantities of zirconia nanoparticles in a stainless steel (316L) powder bed prior to fusion in the LPBF process. The zirconia slurry was deposited as an atomized mist.

The zirconia slurry was deposited as an atomized mist. The deposition process may be adapted to selectively dope a powder bed to form parts with spatially varying properties for use in complex parts. In this work, the added zirconia was shown to be partially distributed in the matrix, but a large portion segregated to grain boundaries and surfaces. Porosity increased in multi-layer parts with the zirconia agglomerated on the porous surfaces. This is undesirable for many applications but may be useful for creating porous surfaces for heat exchanger or bio-integration applications. The deposition of zirconia limits the anisotropic crystal growth throughout the entire doped region.Key words: Laser Powder Bed Fusion, Stainless Steel 316L, ZirconiaIntroductionAdditive Manufacturing (AM), often informally called 3D printing, is attractive to many specialized industries because it is suitable for complex geometry and customized or low volume production [1].

Traditionally, highly sophisticated cutting machines and tools have been used to produce small batches of specialized devices for industries ranging from medicine to aviation [2]. Machining biocompatible metals such as titanium alloys and stainless steel 316L is especially difficult due to work hardening and resulting negative effects on tools and finished products [3]. AM processes, however, allow for cost-effective mass customization by eliminating part-specific tooling. Engineers can also utilize the increased design flexibility of AM to combine multiple parts into single assemblies in a practice called functionality integration [4].Functionally integrated parts can save time and resources by simplifying processing and assembly, reducing material waste, and minimizing weight [5]. A prominent example of functionality integration is a dental strut implant, where an abutment and strut are built as a single part with Laser Powder Bed Fusion (LPBF) [6].

A prominent example of functionality integration is a dental strut implant, where an abutment and strut are built as a single part with Laser Powder Bed Fusion (LPBF) [6]. In this single part, the lower portion must be compatible with bone, while the upper portion must be compatible with human gums. This example highlights one of the greatest challenges of functionality integration. Multi-component1331assemblies frequently require different properties in different parts of the assembly. The dental implant design works well because the property change is only required in one direction. However, functionally integrated designs are much more difficult to execute when an assembly requires different properties in three dimensions [7-9].To overcome this, AM processes need increased spatial control of properties.

Powder Bed Fusion (PBF) is a family of AM processes wherein complex and detailed parts can be built by spreading a thin layer of powder, then fusing a specific cross section of the powder. This process can be repeated thousands of times for large parts. The powders used for PBF include plastics, ceramics, composites, and metals, many of which are engineered for biocompatibility [10-13]. The most common method for metal PBF is LPBF, where the powder is melted by a laser. Binder Jetting (BJ), an alternative to PBF, uses high precision ink jet printing to introduce binding agents to thin layers of powder to form a loosely bound structure, after which this green part is sintered in a furnace [14]. LPBF and BJ both have specific advantages. LPBF can form precise parts without shrinkage and limited need for post-processing. BJ can introduce secondary materials to the part to aid in post-processing or performance [1, 15].

LPBF can form precise parts without shrinkage and limited need for post-processing. BJ can introduce secondary materials to the part to aid in post-processing or performance [1, 15]. A LPBF process that borrows the principle of adding a second material to a powder bed from BJ could make 3-dimensional property manipulation an additional element in the design space of engineers in specialized fields.One of the greatest challenges of changing the properties of parts is choosing secondary materials that can achieve useful property changes with small composition changes [16]. Low carbon stainless steel is a noteworthy material for this study due to its biocompatibility and sensitivity to small changes in composition [17]. Stainless steels are common iron-based alloys that contain chromium and nickel.

Stainless steels are common iron-based alloys that contain chromium and nickel. Carbon and other elements are also frequently added to increase hardness, ductility, and other properties [18].Control of properties and materials is an area of great importance in both AM and biomedical engineering as illustrated by the example given of a dental strut implant. There are many methods available to change properties with small composition changes while maintaining biocompatibility [8, 12, 16, 19-22]. Alloying, carbiding, and oxide dispersion are common methods to modify the properties of metal and have all been used with LPBF. AlMangour et al. increased the yield strength of stainless steel (316L) prepared with LPBF by approximately 35% by pre-mixing steel powder with titanium carbide particles using a ball milling process [23]. This increase in strength is impressive.

This increase in strength is impressive. However, the addition of carbide by this method is time consuming and requires that the entire part be built of the modified material, which would be detrimental if hardness or some property other than strength is required [24].Another notable attempt to modify the properties of LPBF parts uses a high precision vacuum and dispenser system to remove the base material and insert a secondary material in specified locations [7]. Full parts composed of both stainless steel (316L) and copper alloy (Cu10Sn) were built by this method, with each material exhibiting unique properties. This method approaches the flexibility in materials of directed energy deposition (DED), where streams of any material can be directed in front of a high-powered beam and mixed by melting, but with the enhanced environmental control and resolution of LPBF [25, 26]. However, the cost in added fabrication time for multi-powder deposition is significant [15].

However, the cost in added fabrication time for multi-powder deposition is significant [15]. Additionally, depositing the secondary material in this way alters the powder packing, resulting in unpredictable pores, cracks, and other defects where the second material is added [7].1332Oxide Dispersion Strengthened (ODS) Steels are a sub-class of iron-based nanocomposite materials with enhanced properties [27, 28]. ODS parts are typically fabricated from pre-mixed steel/oxide powders and have been processed with LPBF using similar methods to AlMangour, yielding similar results, with similar pitfalls [29, 30]. However, multiple studies show that off-the- shelf oxide nanoparticles can be deposited on a steel powder bed using inkjet technology.

However, multiple studies show that off-the- shelf oxide nanoparticles can be deposited on a steel powder bed using inkjet technology. Davis has explored the influence of various concentrations of Zirconium Oxide (zirconia) on print parameters using single track scans, while Paul demonstrated that yttria nanoparticles deposited on the powder surface can be integrated into the steel matrix, although with some negative affects including porosity due to organic carriers used for deposition [31, 32].This work explores the effects of introducing small quantities (≤ 1 wt%) of zirconia nanoparticles into stainless steel (SS 316L) during the LPBF process for multi-layer specimens to form nanocomposites of biocompatible materials. This study will evaluate the effectiveness of the convective fluid motion in PBF melt pools to disperse zirconia in the metal matrix. Partial re- melting of lower layers in multi-layer specimens is expected to increase oxide mixing [33].

Partial re- melting of lower layers in multi-layer specimens is expected to increase oxide mixing [33]. Additionally, this work will explore the quality of zirconia dispersion in the stainless steel and the impact on the resulting part structure. The change in hardness and the crystalline structure will be compared with unmodified LPBF SS316L. This is the first time that the process will be applied to multi-layer parts.Methods Materials and ProcessingA Concept Laser M2 Cusing dual laser Direct Metal Laser Sintering (DMLS, LPBF) machine was used to form all parts for this study. Three sets of standard laser processing parameters were used to form both single layer and eight-layer rectangular specimens. The standard upskin, downskin, and bulk parameters were used because they have demonstrated good properties for general applications. The specifications for each set of processing parameters include laser spot size, laser scan speed, and laser power (see Table 1).

The specifications for each set of processing parameters include laser spot size, laser scan speed, and laser power (see Table 1). These variables combine to give the energy density, an important factor in predicting part performance. Energy density can be defined in different ways, but for this study a common expression for areal energy density is given by𝐸𝐷𝑎 = 𝑃𝐴 𝑡𝑑 = (𝑃 𝐴𝑠) (𝑑𝑏  𝑣) =𝑃 𝑣𝑑𝑏(4 𝜋)where P is power, td is the exposure time, A is the area covered by the laser, d is laser diameter, v is laser speed. Additionally, 44 single line scans were performed using various input parameters, found in Appendix A. Single line scans provide valuable information relative to the flow of molten material and can be used to predict appropriate inputs for full parts [34]. The ambient atmosphere is nitrogen, with an oxygen content no greater than 0.2%.1333[1]Table 1. Laser input parameters used for area and multi-layer parts.

The ambient atmosphere is nitrogen, with an oxygen content no greater than 0.2%.1333[1]Table 1. Laser input parameters used for area and multi-layer parts. Scan Speed Parameter (mm/s) 980 1350 1350Spot Size (µm)Upskin (US) Downskin (DS) Bulk (BP)130 130 130Laser Power (W) 200 280 370Areal Energy Density (J/mm2) 1.570 1.595 2.108The purpose of this paper is to observe the impact of a zirconia slurry deposited onto a stainless steel powder bed between layers of LPBF process and its impact on the resulting microstructure, (NYACOL and microhardness. The Nanotechnologies Inc.) is a 20 wt% suspension of 100 nm zirconia particles in water with no other additives. The powder is a 316L stainless steel powder from Concept Laser (CL 20ES) with mean particle size of 45 µm.composition,zirconiaslurryThis study is primarily concerned with 2D and 2.5D (extrusions of 2D geometries) parts with uniform doping.

Rather than utilize inkjet printing deposition [31, 32], a piezoelectric atomizer (mister) was used to create a mist of fine droplets, allowing for a greatly simplified deposition process. The mister allows for doping between 3 and 10 mm in a single pass, depending on the distance from the substrate. It forms droplets that spread over a wide area with a relatively uniform distribution, while an inkjet printhead requires precise -x and -y direction controls to cover a similar area, one line at a time (see Figure 1).Figure 1. Large area deposition mode of piezoelectric mister (left) vs small area deposition of inkjet printhead (right).Zirconia deposition was measured by misting onto 5mm x 20mm strips of plastic shim material of known weight. The misting device was attached to a 3-axis stage and moved over a carefully arranged array of shim strips at a constant speed. After drying, the shim strips were weighed to determine the quantity of zirconia deposited per unit area.

After drying, the shim strips were weighed to determine the quantity of zirconia deposited per unit area. The speed at which the device moved over the test area was adjusted until deposition mass was equivalent to 1 wt% of a 0.05 mm thick layer of SS316L in the main column of mist (see Figure 1). The amount of deposited zirconia outside of the central 5 mm below the column of mist was negligibly small.A misting module was designed to integrate with the coater blade of the LPBF machine. A microcontroller was used to precisely trigger zirconia deposition only in the area of interest to avoid contaminating surrounding powder. This simple setup allowed for zirconia deposition in the x direction with the movement of the coater blade, and the z direction with the movement of the build plate.1334The build plate was heated to 70 oC to aid in drying the water from the zirconia slurry.

The LPBF machine was allowed to spread standard layers of 0.05 mm with the misting device depositing zirconia directly after each layer was spread. The process was manually paused for 5 minutes after each layer to ensure that all water had evaporated from the build area, which could cause severe denudation as rapidly expanding water vapor explosively pushed unfused powder away from the laser [35]. The basic process is depicted in Figure 2. Undoped samples were formed with the same laser parameters in parallel to the zirconia doped samples for direct comparison.a abcdeFigure 2. Experimental approach for introduction of zirconia into steel powder bed. a) a flat substrate is prepared to receive a layer of fresh powder. b) A thin layer of powder is spread over the build area by a coater blade. c) The misting device deposits zirconia into the build area directly after fresh powder is spread.

b) A thin layer of powder is spread over the build area by a coater blade. c) The misting device deposits zirconia into the build area directly after fresh powder is spread. d) The powder bed is left to sit for 5 minutes to allow all residual water to evaporate. e) A laser melts the powder, trapping zirconia in the solidified metal matrix.Measurement and CharacterizationAfter laser fusion, the samples were removed from the LPBF machine and cleaned carefully with a toothbrush to remove excess powder. Roughness was measured optically by a Keyence VHX-7000 microscope. The samples were then sectioned via waterjet to obtain a cross- section view of the region where zirconia was deposited. Next, the samples were fixed in a clear epoxy resin and polished to a mirror finish, with 0.05 µm silica vibrational polishing as the final step.The distribution of zirconia in the steel was determined by Energy Dispersive X-ray Spectroscopy (EDS) using a Verios G4 UC Scanning Electron Microscope.

Vickers hardness was measured by a Leco LM100AT microhardness tester conforming to ASTM standard E92-17 with a force of 200 N. Hardness was only measured for the 8-layer samples, where a sufficiently thick layer of steel was fused to the substrate to make hardness indentations. Twenty hardness indentations were performed on each sample in the doped region and at a corresponding location on the undoped samples.The samples were then etched electrolytically in a 10% Oxalic acid solution to expose the microstructure. Optical imaging was performed with the same Keyence microscope used to measure roughness. The image processing software ImageJ [36] was used to analyze these images to measure melt pool dimensions of the single line scans (Figure 3) and determine porosity of the multi-layer parts (Figure 9).1335c100 µmFigure 3. Typical cross section of etched single line melt pool formed with LPBF.

Typical cross section of etched single line melt pool formed with LPBF. The key dimensions of width at the surface plane (a), depth from the surface plane (b), and height from the surface plane (c) are shown.The single line scans were classified by researchers familiar with, but not closely involvedin this work using the classification categories illustrated in Figure 4 as a guide.54325.0 mm1Figure 4. Example single line scans used to aid volunteers in classifying melt pools. Class 5) Continuous melt pool. Class 4) Nearly continuous with some balling. Class 3) Balling across entire scan. Class 2) Balling with some unfused gaps between balls. Class 1) Mostly unfused, but with some balling present.Results and Discussion Single Line ScansBalling is a common defect in LPBF, which can be observed in single line scans. Single line scans with an aerial energy density less than 2.5 J/mm2 were nearly all within the balling regime, as shown in Figure 5.

Single line scans with an aerial energy density less than 2.5 J/mm2 were nearly all within the balling regime, as shown in Figure 5. The addition of 1 wt% zirconia slightly decreased the energy required to achieve continuous or nearly continuous lines. The slight change may be due to the zirconia decreasing the reflectivity of the steel powder, allowing more energy to be absorbed. This supports the findings of Davis, and indicates that the addition of zirconia has little impact on the melt pool-substrate interaction for single lines [31].1336♦ No Fusion ♦ Class 1 ♦ Class 2♦ Class 3 ♦ Class 4 ♦ Class 5BallingContinuousFigure 5. Influence of zirconia on single line fidelity for various energy density inputs. Class 1 includes samples with minimal fusion, and balling where fusion occurs. Class 2 includes samples with increased fusion, but large gaps between balls. Class 3 includes full balling with minimal gaps between balls.

Class 2 includes samples with increased fusion, but large gaps between balls. Class 3 includes full balling with minimal gaps between balls. Class 4 includes nearly continuous melt pools with some balling. Class 5 includes fully continuous melt pools without balling. The continuous and balling regimes are highlighted by colored boxes corresponding to their respective marker colors.Melt pool provides an additional measure of the impact of zirconia on the powder processing characteristics. Melt pool geometries are not easily measured in the balling regime, so only continuous and nearly continuous single line scans were measured. The addition of zirconia generally increased the predictability of melt pool width and depth, as shown in Figure 6, with no significant trends in height from the surface plane compared to undoped melt pools. In contrast, Davis showed an increased melt pool height from the surface plane with the addition of zirconia [31].

In contrast, Davis showed an increased melt pool height from the surface plane with the addition of zirconia [31]. This may be because Davis used inkjet printing of large quantities of dopant and much larger zirconia quantities. These differences may have increased the surface zirconia levels more than the misting methods used here. Both deposition method and concentration could change the laser absorptivity and affinity of zirconia to segregate. In general, melt pool width and depth both increased as energy density increased. This was expected and is supported by the literature [34, 37]. The data contained in Figure 5 and Figure 6 corresponds to similar figures by Davis, and is intended to expand the data sets presented in that work [31].abcFigure 6. Width, depth, and height of single line melt pools.

Characteristic images of the undoped US8 sample and the doped US8 sample are shown. All multi-layer samples have similar appearance.The exceptions to the trend of increased roughness are the single layer downskin and bulk parameter scans, where roughness decreased with zirconia doping. For these areas, it was impossible to clean all the residual zirconia from the surface prior to taking measurements. This residual zirconia likely filled in the deepest cervices, resulting in a smoother surface, but preventing a true measurement of the roughness of only the fused material (Figure 8). To maintain consistency, the roughness of the entire doped region was measured. But a more accurate roughness measurement of only the fused material may have been achieved by more forceful cleaning or by manually measuring regions without agglomerated zirconia and unfused stainless steel powder on the surface.1338bda Abcd1.0 mmFigure 8. Top view of single layer scans.

Top view of single layer scans. a) downskin parameters (DS1) without zirconia. b) downskin parameters (DS1) with zirconia doping. c) bulk parameters (PB1) without zirconia. d) bulk parameters (BP1) with zirconia doping. The lighter grey areas in b) and d) are agglomerated zirconia and steel adhered to the surface.PorosityLPBF can produce parts with near 100% density and good surface finish on exterior surfaces under normal conditions [39]. This holds true for the control group in this study as seen in Figure 9. There was zero porosity measured in any of the control specimens. There was also negligible porosity in the single-layer specimens with zirconia doping. When zirconia was added to multi-layer parts, however, porosity increased to around 12%, with some regions exhibiting up to 30% (Figure 9). The pores are long and nearly continuous.

When zirconia was added to multi-layer parts, however, porosity increased to around 12%, with some regions exhibiting up to 30% (Figure 9). The pores are long and nearly continuous. Similar pores exist in the 8-layer samples with zirconia under all tested process parameters.Figure 9.Zirconia-doped Single layer SS316L specimen with no measurable porosity (top), Undoped 8-layer SS316L test specimen (bottom left), and 8-layer SS316L test specimen with 1 wt% zirconia deposited by the misting technique described for this study (bottom right). All samples etched to reveal grain structure.13390.25mmComposition (EDX)EDX reveals that most of the zirconia agglomerated near porous surfaces within the multi- layer specimens. In Figure 10c, the contrast in the image shows differing atomic weights, where the lightest grey regions are primarily zirconium in the zirconia, the darker grey is primarily iron in the steel, and the black is primarily carbon in the epoxy.

The height of the zirconium peak relative to other peaks in the EDX data (Figure 10a and Figure 10b) indicates the relative quantity of zirconia in the region of interest. Several factors may play into the segregation of zirconia from the stainless steel, and the resulting porosity. Particles on the nano scale have extremely high surface energy, surface electrostatic charge, and van der Walls forces due to their high surface area to volume ratio. These high energy, unstable particles can quickly reduce their total free energy by agglomerating, which is likely why most of the zirconia is found clumped together [40]. aFeCrFecZrNiCrMoCSibZrFeCr FeNiFigure 10. Electron backscatter micrograph with accompanying EDX data. The red and magenta dots on the micrograph indicate the location where the corresponding EDX count was taken.

Electron backscatter micrograph with accompanying EDX data. The red and magenta dots on the micrograph indicate the location where the corresponding EDX count was taken. The arrows indicate x-ray peaks associated with zirconium.Although most of the zirconia agglomerated near pores, Figure 11 indicates that some zirconia dispersed into the stainless steel. Due to the high melting point of zirconia relative to temperatures achieved in this test, it is unlikely that zirconia melted into the stainless steel [41]. Instead, the convective motion of the liquid melt pool likely mixed solid nanoparticles into the bulk stainless steel, resulting in the structure of Figure 11. The concentration is too small to achieve an accurate EDX count. However, the brightness of the circular particles in Figure 11 indicate they primarily consist of an element heavier than iron. Given the composition of SS 316L, it is unlikely that the circular particles are anything other than zirconia.

Given the composition of SS 316L, it is unlikely that the circular particles are anything other than zirconia. The size is also in the expected range of the zirconia particles on the order of 100 nm.1340Figure 11. Backscatter SEM image of stainless steel with zirconia doping. Circular inclusions are likely dispersed zirconia.HardnessThe doped and undoped samples have considerable overlap in hardness as seen in Figure 12. The median hardness of the zirconia doped samples decreased, likely due to porosity adjacent to or below the indentation, as opposed to the undoped samples where no porosity was observed. With outliers removed, the undoped samples were still an average of 27 points harder than the zirconia doped samples on the Vickers scale. While a decrease in hardness is concerning, the average hardness of the zirconia doped samples is still 100 HV harder than as-annealed SS 316L [17].Figure 12. Box and whisker summaries of hardness measurement results.

Box and whisker summaries of hardness measurement results. Outliers, 1.5 IQR or further from the lower quartile, are shown by circles. US, DS, and BP correspond to upskin, downskin, and bulk parameters as given in Table 1.Crystalline structureThe crystalline structure of the zirconia doped samples is starkly different from that of the undoped samples. In Figure 13a the undoped samples show columnar grains elongated in the1341direction of material addition (conventionally the -z direction). In contrast, the zirconia doped specimen shown in Figure 13b shows no pattern of anisotropic grain growth. This may be partially due to pores blocking the growth of crystals. But even away from pores, the anisotropic crystal growth that is characteristic of additive manufacturing is nearly eliminated with the addition of zirconia. This may be caused by enhanced precipitation because of the nanoparticles that mixed into the melt pool [45].

This may be caused by enhanced precipitation because of the nanoparticles that mixed into the melt pool [45]. Isotropic crystal growth mitigates anisotropy in properties, one of the main flaws of AM [15].abFigure 13. Microstructure in multi-layer stainless steel AM specimens. 10a shows a sample with no dopant. 10b shows a sample with a small amount of zirconia deposited on the surface of each layer. Arrows point to several characteristic grains in each image.ConclusionsIn this study, small quantities of zirconia nanoparticles were deposited on the surface of a stainless steel 316L powder bed prior to laser fusion in LPBF. The nanoparticles were evenly distributed across the area of interest in a water suspension by a piezoelectric atomizing mister. The area was then fused by a high-powered laser.

The nanoparticles were evenly distributed across the area of interest in a water suspension by a piezoelectric atomizing mister. The area was then fused by a high-powered laser. The fused steel was measured for roughness, then sectioned and polished to measure the influence of zirconia on single line scans, bulk composition and hardness, and to observe the crystal structure.1. Single line scans indicate that adding small amounts of zirconia does little to change the geometry of the melt pool. There is also a negligible difference between undoped and doped single layer specimens in melt pool size. These results show little influence from the zirconia on the melt pool-substrate interaction. Zirconia in multi-layer parts increased the porosity of the stainless steel. The zirconia mostly agglomerated near the porous interfaces within the stainless steel matrix, similar to the initial results of Paul et al. [32].

The zirconia mostly agglomerated near the porous interfaces within the stainless steel matrix, similar to the initial results of Paul et al. [32]. However, some zirconia mixed into the stainless steel as shown in Figure 11. Re-melting the zirconia- doped region, as per the method of Yu et al. [33] may increase the amount of mixing due to increased fluid motion. However, further experimentation is required to verify this hypothesis.2. The average hardness of the 8-layer stainless steel parts decreased slightly with the inclusion of oxide nanoparticles. This contradicts the results of Zhong et al., Yan et al., and Pawawoi et al. [27-29]. The contradiction may be explained by the porous test surface, shown in Figure 9, which reduces the reliability of microhardness testing. A non-porous1342test surface may be achieved by changing laser scan parameters, especially increasing areal energy density [39, 46].3.

A non-porous1342test surface may be achieved by changing laser scan parameters, especially increasing areal energy density [39, 46].3. Although the mechanical properties of the stainless steel were reduced, anisotropy in the crystal structure was nearly eliminated (Figure 13). LPBF generally forms parts with dendritic crystals growing in the direction of material addition. The minimization of dendritic crystals typically results in a stronger, but a less ductile, final product. A similar result is expected for LPBF if full density parts are achieved without dendritic crystal growth. In this case, anisotropic crystal growth was likely reduced due to zirconia nanoparticles serving as nucleation points, making nucleation of new crystals more energetically favorable than growth of existing crystals [45, 47].4. The porosity of the zirconia-treated samples is a critical limitation for most LPBF applications.

The porosity of the zirconia-treated samples is a critical limitation for most LPBF applications. However, this may be favorable as a surface treatment in some biological applications. Porosity can be beneficial for osseointegration of implants, and is often a sought-after feature of medical devices, especially those formed with AM [11, 48-50]. The porous surface of Figure 9 resembles the bone scaffold structure of Jiao et al [51], and the porous surface structures of Wu et al [13] and Kwak et al [52] intended for osseointegration. Additionally, zirconia is a biocompatible material that has been shown to behave similarly to other oxides used for ODS strengthening [42], which improve corrosion resistance even in the most extreme environments [53].

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceImpact Strength of 3D Printed Polyether-ether-ketone (PEEK)Haijun Gong1,*, Xiaodong Xing2, Jan Nel11Department of Manufacturing Engineering, Georgia Southern University, Statesboro GA, USA2College of Mechanical and Electrical Engineering, Harbin Engineering University, Harbin, ChinaCorresponding author (hgong@georgiasouthern.edu)AbstractPolyether-ether-ketone (PEEK) is a high-performance thermoplastic with high heat-, high chemical-, high water-, and high wear-resistance. Its strength and durability also make it highly accepted for a range of industrial applications. 3D printing of PEEK filaments offers a unique quality and flexibility in making PEEK parts for low-volume production or special designs. This study investigates the impact strength of 3D printed PEEK materials.

This study investigates the impact strength of 3D printed PEEK materials. The specimens are fabricated using a fused deposition modeling (FDM) based 3D printer and tested by a pendulum impact tester in compliance with ASTM standard. The testing result is discussed with respect to the processing parameters and the annealing treatment. Impact strength comparison of PEEK materials manufactured by 3D printing and by conventional production is also conducted.InstructionPolyether-ether-ketone (PEEK) is a semi-crystalline polymer, which was developed in 1978. The PEEK is widely used as a high-performance thermoplastic due to its high heat-, high chemical-, high water-, high wear-resistance. Especially, PEEK and its varieties have been adopted for different biomedical treatment such as dental implants, orthopedics and maxillofacial surgery [1,2]. Its strength and durability also make it highly accepted for a range of industrial applications.

Its strength and durability also make it highly accepted for a range of industrial applications. For example, many automotive applications require to work at a temperature more than 120 °C. A high continuous operating temperature (260 °C) and excellent mechanical properties at this temperature make PEEK polymer the natural choice [3]. 3D printing of PEEK filaments offers a unique quality and flexibility in making PEEK parts for low-volume production or special designs. The mechanical properties of 3D printed PEEK material are of great interests to the additive manufacturing research community.Rinaldi et al. [4] conducted mechanical, thermal (DSC), microstructural (XRD), and morphological testing of FDM printed PEEK samples. The results outlined the impact of printing parameters on the final mechanical performance and internal porosity. But there is no thermal and microstructural difference between PEEK filament and part. Want et al.

But there is no thermal and microstructural difference between PEEK filament and part. Want et al. [5] used finite element analysis (FEA) to simulate the melting conditions and fluidity of PEEK in a flow channel, in order to determine the parameters required to 3D print PEEK parts with sufficient surface quality and improved mechanical properties. The validation experiments suggested the heating temperature and printing speed, as well as layer thickness, to ensure a desired density and strength with minimized defect. PEEK can also be produced with carbon nanotubes (CNTs) in the FDM system to make CNT PEEK composites [6]. The influence of CNT on the mechanical performance and715microstructures was studied. The effect of thermal processing conditions on crystallinity and mechanical properties on 3D printing PEEK was studied by Yang et al. [7].

The effect of thermal processing conditions on crystallinity and mechanical properties on 3D printing PEEK was studied by Yang et al. [7]. The results indicate that the temperature has a critical impact on the degree of crystallinity and mechanical properties of PEEK parts.However, there is few researches about the impact strength evaluation of 3D printed PEEK material. The impact test measures the toughness of a material. Although the material toughness is not directly used for design purpose, it indicates the ability of the materials to withstand against impact load. The impact test is useful for understanding the resistance of the different materials or the same material in different processing conditions to impact loading. The purpose of this research is to study the impact strength of 3D printed PEEK material.ExperimentalCreatBot F430 high-temperature 3D printer is used for fabricating the PEEK impact testing specimens.

It has two hotends, one of which can ben heated to 420 °C to print PEEK filament, as shown in Fig. 1. The PEEK filament (non-medical grade) came with the 3D printer, with a diameter of 1.75 mm.High-temp hotendFig. 1 High-temperature 3D printer and hotendsImpact testing specimen has a standard size for Izod pendulum impact resistance test, conforming to ASTM D256. A hammer is raised to a specific height and then released. The hammer swings down hitting a notched specimen. The energy absorbed by the specimen is measured to calculate the impact strength. The 3D printer exerts printing for a 100% fill rate to the infill region (raster scan) after double contouring scans, as illustrated in Fig. 2. The direction of infill scan is re- oriented 90° every layer to ensure a good bonding. The notch of the specimen is created directly by 3D printing, instead of being cut by a notcher.

2. The direction of infill scan is re- oriented 90° every layer to ensure a good bonding. The notch of the specimen is created directly by 3D printing, instead of being cut by a notcher. The purpose is to see if the “printed notch” similarly causes a dramatic stress concentration and crack propagation resulted from impact load. An Instron CEAST 9050 pendulum impact tester (Fig. 3) is employed to measure the Izod impact strength. PEEK is a semicrystalline thermoplastic, with a glass transition temperature of 143 °C (289 °F) and melting temperature of 343 °C (662 °F). The heating temperature needs to be elevated higher than the melting point to ensure a continuous extrusion. The melted PEEK exhibits a fast solidification rate from liquid phase to solid. Also, the scan speed must be slow for PEEK filament, as suggested by the machine vendor. The process parameters of 3D printing PEEK are listed in716Table 1.

Also, the scan speed must be slow for PEEK filament, as suggested by the machine vendor. The process parameters of 3D printing PEEK are listed in716Table 1. Annealing is conducted in the furnace to the impact specimens to improve the crystallinity. The annealing recipe is shown Fig. 4.Fig. 2 Izod specimen and scan pattern Fig. 3 Instron CEAST 9050 impact testerTable 1. 3D printing PEEK parametersLayer thickness (mm)0.15TAnnealing150 "C, 3 hFill density (%)100Scan speed (mm/s)20Nozzle temperature (°C)420Bed temperature (°C)80Fig. 4 Annealing temperature of 3D printed PEEKResults and DiscussionDeformation and Delamination of PEEK Parts. The 3D printed PEEK specimens exhibit apparent deformation (warping) due to the thermal strain as shown in Fig. 5(a). The warping deformation has been proved to be closely correlated to the chamber temperature and nozzle temperature [8].

5(a). The warping deformation has been proved to be closely correlated to the chamber temperature and nozzle temperature [8]. The deformation results in delamination of PEEK specimen off the build plate, which might cause serious contacting with printing head. Otherwise, the as-built PEEK specimen (non-annealed) presents a weak bonding strength of between layers, as shown in Fig. 5(b). The crack can be easily developed when load is applied to the specimen for removal from the build plate. Extra caution should be used to take the specimen off the plate.CrackDeformationFig. 5 Deformation and interlayer crack of 3D printed PEEK Parts717Annealing and crystallinity. The crystallinity levels of the as-built PEEK specimens are increased through the annealing heat treatment. After annealing, the thermal history, dimensional changes, and residual stress are removed, correspondingly resulting in an improved mechanical property.

After annealing, the thermal history, dimensional changes, and residual stress are removed, correspondingly resulting in an improved mechanical property. The enhanced bonding strength between layers is a strong evidence of the improvement. A comparison of impact testing to the as-build specimen and the annealed specimen also demonstrates the effect of annealing to the PEEK parts. As shown in Fig. 6, when the hammer hits the as-built PEEK specimen, the impact load does not cause a crack initiated from the notch. Instead, it causes interlayer deboning of the rear face (Fig. 6(a)). The weak bonding is not desired to the 3D printed PEEK parts, since it fails to resist the impact loading as expected. On the contrary, the annealed PEEK specimen shows an expected failure mode, which has a crack initiated from the notch. The rear face of the specimen (Fig.

On the contrary, the annealed PEEK specimen shows an expected failure mode, which has a crack initiated from the notch. The rear face of the specimen (Fig. 6(c)) exhibits good bonding between layers after the specimen is broken.(a)(b)Rear faceAs-built specimen (no annealing)(c)(d)Annealed specimenRear faceFig. 6 Comparison of as-built and annealed specimens after impact loadingImpact strength. The impact strength of 3D printed PEEK specimens is shown in Fig. 7. Some reference values are included in the same diagram for comparison. It is noted that the impact strength (73±12 KJ/m2) of the PEEK specimens tested in this study is much higher than other notched PEEK materials. The value is comparable to the un-notched Victrex PEEK 450G material (impact strength 70 KJ/m2) which is glass fiber reinforced. The high impact strength value deserves a further confirmation with more tests.

The high impact strength value deserves a further confirmation with more tests. Otherwise, the impact strength value also indicates the 3D printed notch does not perform the same to the notch cut by a notcher. A comparison study is necessary for the future research. In addition, the annealing recipe is believed to influence the material properties in a certain extent. The crystallization of polymer is associated with the partial alignment of its molecular chain. The impact strength of the semi-crystalline PEEK material is718determined by the crystallinity and the size/orientation of the molecular chains.

The impact strength of the semi-crystalline PEEK material is718determined by the crystallinity and the size/orientation of the molecular chains. Both annealing and scan strategy play an important role to the impact property of 3D printed PEEK material.)8073±12702m / J K( h t g n e r t S t c a p mI604020Data source: (cid:120) (cid:120) (cid:120)www.victrex.com www.polyfluor.nl/en/materials/peek/ www.intamsys.com/product/intamsys- peek-3d-filament/8105190PEEK (tested in this study)Victrex PEEK 450G (notched)Victrex PEEK 450G (glass fiber reinforced, notched)Victrex PEEK 450G (glass fiber reinforced, un- notched)Polyfluor PEEK filament (notched)Intamsys PEEK filament (notched)Fig. 7 Impact strength of 3D printed PEEK materialsFractography. The fracture face of the PEEK specimen is shown in Fig. 8. It is noted that the fracture is primarily attributed to the tension rupture of the extruded thin strands.

The fracture face of the PEEK specimen is shown in Fig. 8. It is noted that the fracture is primarily attributed to the tension rupture of the extruded thin strands. The bonding force between layers or between strands also contributes to the impact strength. But the accumulated tension strength of the thin strands is of great importance to the impact loading performance, due to the aligned molecular chains.Fig. 8 Fracture face of the 3D printed PEEK impact specimen719ConclusionThis study concentrates on the impact testing (Izod pendulum testing) of 3D printed PEEK material. It shows that the annealing process improves the levels of crystallinity of PEEK specimens, which has a high impact strength. More tests deserve to be conducted to validate the result. Also, the 3D printed notch is not the same as the cut notch. The accumulated tension strength of the extruded thin strands contributes much to the impact strength.Reference[1] S. Singh, C. Prakash, S.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceImpact Testing of 3D Printed Kevlar-Reinforced Onyx MaterialMackenzie Scrocco1 ​, Timothy Chamberlain1 1 1 ​, Edward Ellks ​, Logan Weinreber ​, Chloe Chow​​​​1 2 ​, Brett P. Conner ​, Pedro Cortes Connor Halford1​​​1 ​,​1 ​Youngstown State University 2 ​Youngstown Business IncubatorAbstract:Kevlar is a common reinforcement used in composites for applications requiring impact resistance. This research evaluated the impact performance of 3D printed continuously reinforced Kevlar fiber on 3D printed nylon composites containing chopped carbon fiber (CCF). The 3D printer used was a Markforged Mark Two material extrusion system, and Onyx is the trade name of Markforged’s nylon/CCF material.

The 3D printer used was a Markforged Mark Two material extrusion system, and Onyx is the trade name of Markforged’s nylon/CCF material. In this work, a diverse number of composite architecture designs including type of infill pattern, number of Kevlar layers, and location of Kevlar layers were investigated. The printed specimens were characterized using a Charpy impact testing and gas gun ballistic testing. Results were compared on the basis of weight and estimated part cost.Keywords:impact testing, kevlar, Markforged, projectile​1. IntroductionThe impact resistance of materials is needed for a variety of applications. In aviation, gas turbine engine containment rings prevent detached rotating components from exiting the engine . Forward aircraft structures and cockpit windows need to nacelle and impacting the aircraft [1] . Manned space vehicles require a level of protection resist bird strike and hail impact [2–4] against micrometeoroid impact.

Manned space vehicles require a level of protection resist bird strike and hail impact [2–4] against micrometeoroid impact . Similarly, high performance materials are desired for body [5] and vehicle armor to stop projectiles while minimizing weight​​​​​​[6,7].​​Additive manufacturing (AM) provides advantages over conventional manufacturing processes by providing a more capable and cost effective means to produce complex geometries . Indeed, AM yielding more functional designs, and an efficient means of customization allows for mass customization, production on demand or production at the point of need, and quantity flexibility leading to affordable low volume production and the ability to shift into[8]​​1121higher production quantities on the same equipment . These advantages can be brought into the design and fabrication of impact resistant structures.

These advantages can be brought into the design and fabrication of impact resistant structures. A key first step is understanding the mechanical performance of these 3D printed structures.Often the optimal combination of impact resistance and weight come from multi-material solutions . Therefore, multi-material additive manufacturing solutions should be of interest to create lightweight-high performance systems. One approach is to provide fiber reinforcement to polymeric parts. The Markforged Mark Two shown in Figure 1 is a desktop material extrusion 3D printer. The Mark Two is capable of printing a filament consisting of nylon and chopped carbon fiber (CCF). This material is called Onyx and is 1.4 times stronger and stiffer than ABS [12] . The Mark Two also has a second extruder which allows for a continuous fiber reinforcement within the printed part. The Mark Two is capable of printing the following materials as fiber reinforcement: Carbon Fiber, Fiberglass, Kevlar.

The Mark Two is capable of printing the following materials as fiber reinforcement: Carbon Fiber, Fiberglass, Kevlar. Markforged claims the Mark Two has a 10 micron accuracy meaning that a build could be paused, the build plate removed, and objects such as threaded nuts or bolts could be incorporated into the printed parts and the build plate brought back into the build envelope for continued printing. This feature was not verified in this study.[9–11]​​​Figure 1: The Markforged Mark Two printer is shown at left. At right the two extruder nozzles are shown. The left nozzle in the image is for the continuous fiber reinforcement and the fiber filament spool is visible. The Onyx nozzle is at the right of the image.The Markforged Eiger software is used for build processing including file slicing, layer heights, infill pattern selection, reinforcement pattern selection, location of reinforcement, and location of parts on the build plate.

The Onyx layer heights that can be selected for unreinforced parts are: 0.100, 0.125, and 0.200 mm. As shown in Figure 2, several Onyx infill patterns can be selected: triangular (TRI), rectangular (RECT), and hexagonal (HEX). There are also options for reinforcement infill patterns as shown in Figure 3. The first reinforcement pattern is isotropic1122(ISO) which is similar to the RECT infill pattern. The second fiber reinforcement pattern is concentric (CON), which prints concentric rings of fiber around the perimeter of the part. ISO and CON can be combined in the same layer. If reinforcement is selected, the Onyx layer height is restricted to only 0.100 mm. Combinations of these parameters can be used to create composite structure architectures.Figure 2: Onyx infill patterns​Figure 3: Continuous fiber reinforcement fill patterns. Top: isotropic, concentric.

Top: isotropic, concentric. Bottom: Charpy v-notch specimen with combined isotropic and concentric reinforcement.1123The effect on tensile properties of continuous fiber reinforcement of 3D printed parts was evaluated by Melenka, et al . In their study, the researchers used a Markforged Mark One printer. The base material was nylon not Onyx, and the continuous reinforcement material selected was Kevlar. The continuous fiber fill parameter was chosen, and specimens with zero reinforcement rings (0% volume fraction), two rings (4.04% volume fraction), four rings (8.08% volume fraction), and five rings (10.1% volume fraction) were evaluated. The researchers found that the tensile modulus of the nylon was on the order of 1 MPa and the ultimate tensile strength was less than 10 MPa. The researchers experimentally determined that the Kevlar reinforcement resulted in dramatic increases in the elastic moduli and the tensile strength.

The researchers experimentally determined that the Kevlar reinforcement resulted in dramatic increases in the elastic moduli and the tensile strength. For the two rings, the modulus was 1767.2 MPa and the strength was near 30 MPa. In contrast, for the four rings, the modulus was 6920.0 MPa with a strength near 60 MPa; whereas for the five rings, the modulus was 9001.2 MPa and the measured strength was near 90 MPa.[13]​​Caminero et al examined the effect of carbon fiber, Kevlar, and glass fiber reinforcement on nylon (not Onyx) using 3D printed Charpy v-notch specimens manufactured on a Markforged Mark Two printer . The architecture chosen involved a rectangular fill of maximum density with isotropic reinforcement. Three process parameters were examined: layer thickness of unreinforced samples, build orientation, and fiber content. The researchers printed Charpy specimens following the ASTM D6110 specification, and these were printed in two orientations.

The researchers printed Charpy specimens following the ASTM D6110 specification, and these were printed in two orientations. The first was called the “Flat” direction where the specimen lies flat on the XY plane of the build plate and the notch in the X direction (implying a fracture propagation in the X orientation with the notch thickness being in the Z direction). The second orientation was referred to as the “On-edge” direction where the notch faced in the Z direction and therefore fracture propagated in the Z-direction. They concluded that impact strength increased as layer thickness increased for the flat orientation. However, impact strength decreased in the on-edge direction. By printing flat and on-edge specimens with reinforcement layers being oriented in the XY plane of the build, the researchers evaluated the effects of orientation and reinforcement. On-edge samples exhibited higher impact strengths than flat samples.

On-edge samples exhibited higher impact strengths than flat samples. Impact strength was also shown to have increased with fiber volume content. Finally, glass fiber showed the best impact performance followed by Kevlar and lastly the carbon fiber reinforcement[14]​​[14]​.​2. MethodologyIn this work, the Onyx without reinforcement, and the Kevlar reinforced Onyx composite were 3D printed and then mechanically evaluated under dynamic impact conditions. Here, two forms of testing were used to characterize architectures. The first was a Charpy v-notch impact testing, and the second a gas gun projectile testing.1124The Charpy v-notch testing was carried out in order to screen composite architectures for impact resistance, since Charpy specimens are inexpensive as compared to the gas gun targets. Energy absorbed, cost, and weight were the evaluation criteria for this testing.

Energy absorbed, cost, and weight were the evaluation criteria for this testing. It is worth mentioning, that the Charpy v-notch testing equipment at YSU is for metals testing (ASTM E23). As a result, the Charpy v-notch specimen geometry from ASTM E23 was used. The specimen dimensions were 10 mm x 10 mm x 55 mm with the notch edges each being 2 mm long at a 45 degree angle from each other with a fillet radius of 0.25 mm. If absolute values were of interest instead of comparative, ASTM D6110 – 18 would be more suitable. Twenty architectures were designed for the Charpy testing. The architectures are defined in Table 1. Layer thickness, Onyx infill, and Kevlar reinforcement parameters were here investigated.Specimen orientations are shown in Figure 4. “XY Notch Down” Charpy specimens had the length and width in the XY build plane but the notch was face down.

“XY Notch Down” Charpy specimens had the length and width in the XY build plane but the notch was face down. “XY Notch Side” Charpy specimens had the length and width in the XY build plane but the notch was oriented in the side direction. Z oriented specimens had the length parallel to the Z axis. For comparison, , while XY Notch Side XY Notch Down corresponds to the on-edge orientation from corresponds to flat. All reinforcements were printed in the XY plane.[14]​​The YSU Gas Gun Launcher, by REL Inc., shown in Figure 5 is used to accelerate projectiles at chosen pressures for testing the impact strength of materials. Projectiles are held by a plastic sabot. The cylindrical sabot is placed in a pressured chamber, which release the projectile through a pressure differential created by opening a release valve. Once the projectile is fired, it travels through piping until it hits the target securely mounted in a trunk-frame.

Once the projectile is fired, it travels through piping until it hits the target securely mounted in a trunk-frame. The gas-gun contains a chronograph that measures the projectile velocity. Here, the impact energy is evaluated by using the kinetic energy. For this gas gun testing, McMaster Carr hardened 440 stainless steel spherical bearings were used as the projectiles. These bearings have an 8 mm diameter and were measured to have an average mass of 2.0550 g. The projectile diameter was greater than the thickness of the plates (101 mm x 101 mm x 6 mm). For most tests, a 100 psi pressure was selected which normally corresponds to a velocity of on the order of 250 m/s. The sabots consisted of 3D printed ABS plastic, manufactured on Lulzbot TAZ 6 and CreatorBot3D desktop printers. The sabots were 25 mm in diameter and 25.5 mm tall, and contained an 8mm center hole, where the steel bearing was placed.

The sabots were 25 mm in diameter and 25.5 mm tall, and contained an 8mm center hole, where the steel bearing was placed. It is worth mentioning that the targets used in gas gun projectile testing needed considerably more material than the Charpy specimens, and therefore, were relatively expensive.Both Charpy and gas gun specimens were optically characterized using digital photography and low magnification microscopy using a Keyence VHX-6000. Cross-sectioning of gas gun specimens was accomplished using a band saw.1125Architect.

Results and discussion3.1 Charpy Impact TestsThe results of the Charpy Impact tests are shown in Table 2 along with the mass and costfor each architectures. All cases had 3 to 4 replicates.Comparing the impact performance between the XY Notch Down and Z orientation samples (for reference see figure 4), it was observed that the Z orientation is noticeably weaker regardless of the Onyx infill pattern type examined (TRI or RECT). For instance, the reduction in impact strength from Specimen #6 to a similar triangular infill but Z orientation Architecture #4 is 42%. The reduction in impact strength from Architecture #16 (XY notch down) to a similar rectangular infill based on Z orientation such as Architecture #13 is 62%. Here, the comparison between a non-Kevlar reinforced specimen (sample #16) and a Kevlar reinforced Z specimen (sample #13) is valid because the Kevlar is a non-factor in Z orientation, as will later be discussed.

The best performing architectures in the Z orientation were the rectangular Onyx infill. For example, compare sample #10 (TRI) against specimen #13 (RECT).) did not see a significant difference in impact strength between layer thickness of 0.100 mm and 0.200 mm (compare Architecture #2 and #7). This observation differs from that of Caminero et al that showed an increase in impact strength for XY Notch Side as layer thickness increased. For XY Notch Down (or “On-Edge” in ), the impact strength of 0.200 mm is greater than that of 0.100 mm and that of 0.125 mm. This appears to contract findings reported by Caminero et al for XY Notch Down (or “On-Edge”). This difference may be the result of the Onyx being more brittle than nylon, a performance associated with the presence of the chopped fiber reinforcement within Onyx.For the effect of layer thickness, XY Notch Side (a.k.a.

“flat” in[14]​​[14]​​Unsurprisingly, Kevlar provides no useful role on the impact performance of samples containing the notch parallel to the Kevlar planes; as is the case of the reinforced Z orientation specimens (refer to architecture #3 and #10). However, when the Kevlar reinforcement layer was perpendicular to the notch plane as would be the case in the XY Notch Down orientation, then more energy was required to fracture the specimen. This is observed on architecture #1, which 3 does not contain Kevlar, and sample #8, which has the same architecture but contains 0.27 cm of Kevlar in front of the notch and at the rear of the notch. Here, there was an increase of 47% of the energy required for fracture the specimen with the addition of Kevlar.

Here, there was an increase of 47% of the energy required for fracture the specimen with the addition of Kevlar. There was an even more dramatic difference when comparing architecture #16 to #15.In this case, by adding a 3 of Kevlar to the architecture of architecture #15, the impact energy monolithic layer of 1.49 cm1128increased from 0.832 lb ft to 8.394 ​lb ft. Indeed, the reinforcement being parallel to the XY plane, requires the crack to penetrate the entire layer. In practice, as shown in Figure 6, the fracture path becomes exceptionally tortuous as the crack attempts to progress by delaminating the reinforcement layer away from the Onyx, and tearing the fibers when possible. It should also be noted that the non Kevlar reinforced rectangular Onyx infill specimen (#16) outperformed the non Kevlar reinforced triangular specimen (#1).

It should also be noted that the non Kevlar reinforced rectangular Onyx infill specimen (#16) outperformed the non Kevlar reinforced triangular specimen (#1). The rectangular infill pattern has a higher fill setting while the triangular has a narrower infill band.In the case of the XY Notch Side orientation, the addition of Kevlar slightly increases the impact energy of the samples. Comparing architecture #2 which does not contain Kevlar to #9 3 which has the same architecture but with 0.26cm ​of Kevlar at the top and bottom of the specimen, it was seen that the presence of the reinforcement increased the impact energy of the printed system by 11%. In the XY Notch Side orientation, only the edge of the reinforcement layer is presented to the crack and tearing and bypassing of the Kevlar is easier at this orientation.​While the highest performance was achieved by Kevlar reinforced specimens with RECT Onyx infill, both cost and machine time are high for such a combination.

A good mix of cost, weight, and performance was obtained by HEX Onyx infill patterns with combined ISO and CON reinforcement fill patterns (see architectures #18 to #23).1129Figure 6: Images of specimens of architecture 15 which is reinforced, versus unreinforced​architecture 16. The micrographs show distinctive differences in their fracture behavior, due to presence of the Kevlar fibers.

Measured Charpy Impact energy along with calculated mass and cost from Eiger softwareTable 2:​Specimen NameAverage Energy (lb ft)Total Mass (g)Cost10.6244.72$0.9420.7134.72$0.9530.2804.8$0.9640.4464.71$0.9450.4714.66$0.9360.7644.66$0.9370.7264.66$0.9380.9185.83$1.6390.7905.84$1.62100.25410.57$2.18111.4075.77$1.45120.43412.59$2.58130.31912.59$2.58140.9907.8$2.01158.3947.58$5.69160.8337.01$1.40173.4827.01$1.40183.5696.51$2.49193.1536.36$2.52201.5006.16$1.81212.7896.16$1.81223.7866.34$2.46232.1346.35$2.46244.0006.51$2.491130Figure 7: Mass-impact energy relationship of the composites printed by a MarkForged system.​Figure 8: Specimen material cost is plotted against average impact energy​1131The Charpy test results were here used as an architectural design guidance for the gas gun testing.

Figure 7 shows a plot of the average impact energy versus mass, while Figure 8 is a plot of the average impact energy versus cost. The addition of Kevlar is necessary. Higher loadings of Kevlar greatly improved performance but architecture must be taken into account. Architecture #20 had a 1 mm layer of Kevlar at the back side of the specimen away from the notch while #23 doubled the layer at the back side of the specimen. Neither did well but that is likely a function of the Charpy geometry -- by the time the crack reached the back end the specimen would have dramatically folded over. This might not be as relevant in the gas gun test, so it should be tested. But it does indicate architecture matters. Rectangular infill with high density provided high average impact energies but also additional costs.3.2 Gas gun testsFour architectures were designed based on the results of the Charpy impact tests. These are shown in Table 3. Architecture G1 had no reinforcement.

These are shown in Table 3. Architecture G1 had no reinforcement. G2 had reinforcement at one face of the plate. G2 and G3 had nearly the same percent volume of Kevlar reinforcement (36% and 37%) but the Kevlar was placed at two different locations in the samples. Adding reinforcement increases build time due to the physical motion of a second extruder head and the differences in deposition speeds. Kevlar is also much more expensive than Onyx, so the price is greater.All of the samples here were 6 mm thick plates and all gas gun tests performed were an overmatch. In other words, the projectile completely penetrated the plate for each test. Even reducing the velocity from around 255 m/s to 210 m/s (by reducing the gas pressure from 100 psi to 50 psi) still produced an overmatch on a reinforced plate. Many more tests and plates would have been required to determine a V50 velocity for each architecture at this thickness.

Many more tests and plates would have been required to determine a V50 velocity for each architecture at this thickness. However, based on the fracture mechanisms shown by the perforated samples, a performance impact evaluation was here performed. The results are summarized in Table 4.For the G1 plate architecture, two plates were printed and tested at 255 and 208 m/s. As previously mentioned, both tests completely penetrated the plate as shown in Figure 9. From the sample tested at 255 m/s, it is observed that the back face (opposite of the impact site) shows a damage zone of nearly 14 mm of diameter, with a distinctive X-shaped delamination form. One delamination petal was torn away from the target during projectile exit revealing the triangular infill pattern below it.

One delamination petal was torn away from the target during projectile exit revealing the triangular infill pattern below it. The X shape of the delamination is likely due to the orientation and rotation of the raster pattern in the face sheets of the build during the building process.1132Table 3:Architectures for gas gun testing​Architect. NumberLayer thickness (mm)Onyx InfillKevlar reinforcement % Vol KevlarMaterial CostEstimated Print TimeG10.100TRINo0$9.535 h 38 mG20.100TRI5 mm Onyx followed by 0.5 mm Kevlar then 0.5 mm Onyx8.44$17.36 h 55 mG30.100HEX0.5 mm Onyx on bottom followed by 1 mm of Kevlar, 3 mm of Onyx, 1 mm of Kevlar, 0.5 mm of onyx on top35.94$45.7410 h 11mG40.100HEX2 mm thick Kevlar ISO and CON reinforcement at center37.06$47.6313h 2mTable 4:Summary of gas gun results​Architect.

NumberShot #Measured Plate Mass (g)Pressure (psi)Round velocity (m/s)Result (CP=complete penetration)G1#1 (Plate 1) #2 (Plate 2)Plate 1: 42.84 Plate 2: 42.92100 50255.0 208.4CP, small X-shaped exit delamination CP, large X-shaped exit delaminationG2#1 (Plate 1) #2 (Plate 1) #3 (Plate 1) #4 (Plate 2)Plate 1: 46.77Plate 2: 46.85100 100 100 100No Data No Data No Data No DataFor Shots #1-3 CP, small hexagonal shaped delamination and complete removal of the petals. Shot #4: CP, small X-shaped exit delaminationG3#1#257.14100100256.5264.4CP, left side of target, large X-shaped exit delamination CP, center of target, large X-shaped exit delamination, large X-shaped entrance delaminationG4#158.69100257.6CP, large X-shaped exit delamination, bulge around entranceFor the lower velocity (208 m/s) impact, the X-shaped delamination zone was nearly 30 mm a side.

This exit damage zone is nearly 5X times that of the zone in the 255 m/s test. This suggests that if the velocity is high enough, the projectile will pass through the target with minimal resistance leaving a small exit hole. At lower velocities such as 208 m/s, the printed1133samples would be able to catch the projectile. Here, the structure will “bulge”, leaving a zone of major plastic deformation around the exit hole.Figure 9: G1 Plate 1 at left and G1 Plate 2 at right​The G2 architecture was reinforced only on one side of the specimen. Here, depending on how the plate was oriented with respect to the incoming projectile, the Kevlar might be closer or further from the impact point.

Here, depending on how the plate was oriented with respect to the incoming projectile, the Kevlar might be closer or further from the impact point. This configuration was tested at 100 psi, which represents a velocity close to 250 m/s; however, error in the chronograph prevented the capture of the impact velocity, and consequently the impact energy.G2 Plate 1 was shot three times in the orientation with the reinforcement at the front face. Plate 1 is shown in Figure 10. Two shots were relatively close to each other such that the exit damage zones overlapped. Figure 11 is a photograph of a cross section of the damage zone of the two neighboring impacts while Figure 12 is a higher magnification image of one of the holes taken from the microscope. There is considerable ductile plastic deformation of the Kevlar fiber layers. Followed by a conical damage zone of spall that delaminates the back wall of the target plate.

There is considerable ductile plastic deformation of the Kevlar fiber layers. Followed by a conical damage zone of spall that delaminates the back wall of the target plate. The back wall is completely separated from the target (evident on all exit holes in Figure 10) but the damage zone is relatively small meaning that this architecture is highly overmatched.Figure 13 is a microscope image of the stand-alone hole in G2 Plate 1. It should be noted that entry hole is smaller than the projectile diameter. This indicates that there was expansion to allow the projectile to pass and followed by elastic contraction. It is elucidated, that the1134triangular infill has a degree of compliance that accommodates the projectile while passing through the sample. The exit zone shape is characteristic of all of the G2 Plate 1 impact sites in that it is hexagonal.

The exit zone shape is characteristic of all of the G2 Plate 1 impact sites in that it is hexagonal. This is likely due to the triangular infill.Figure 10: G2 Plate 1 front face at left and back face at right.​Figure 11: G2 Plate 1 cross section​1135Figure 12: Close up image of the right hole seen in Figure 11.​Figure 13: Impacted G2 Plate 1 based on a triangular infill. Entry hole (left), and exit zone​(right).The impacted G2 Plate 2 with the Kevlar placed in the back of plate is shown in Figure 14. At the exit damage zone, the back face wall is delaminated but not torn away. There is more containment than that observed in Plate 1, but the architecture is still highly overmatched as evidenced by the small size of the exit zone. Figure 15 shows the cross section of the G2 plate 2. In this figure, it is shown that the spall region is narrower and more cylindrical than the displayed in plate 1.

Figure 15 shows the cross section of the G2 plate 2. In this figure, it is shown that the spall region is narrower and more cylindrical than the displayed in plate 1. Here, a clear Kevlar delamination layer extending well beyond the exit hole is observed, but no spall cone was observed.1136Figure 14: Impacted G2 Plate 2. Front face (left) and back face (right).​Figure 15: Micrograph of G2 Plate 2. The front face is on the top of the sample.​For architecture G3, the reinforcement was symmetric so orientation was not an issue. Two shots were fired. The first was left of the target center due to debris in the sabot catching system resulting in a skewed shot. However, there was a large half X shaped delamination (the half X due to the proximity to the target edge which was fixed). This offered an opportunity for a the center. The G3 target plate is shown in Figure 16.

This offered an opportunity for a the center. The G3 target plate is shown in Figure 16. second shot which did impact Interestingly, a massive delamination zone appeared at the point as well as at the exit point. ​ The large exit damage zones infers that this architecture is has a superior impact performance compared to G1 and G2 due to the impact energy being dissipated in large delamination areas. If the plate thickness can be increased while maintaining the percent volume infill, this architectureentry​1137would stop the projectile at this velocity. The large entry site delamination is possibly the result of a shockwave deflection along the front face or due to a spring-back effect from the near-surface Kevlar layer after the projectile has passed. There may have also been an accumulation of damage from the first shot.Figure 16: Impacted G3 target plate. Front face (left), and back face (right).​Figure 17 is a cross sectional view of the G3 target plate.

Front face (left), and back face (right).​Figure 17 is a cross sectional view of the G3 target plate. Both impact sites are shown.There is significant plastic deformation of the Kevlar layer and the Onyx layer sandwiched between. Each Kevlar layer in this architecture is twice as thick as the Kevlar in G2. A high magnification image of the center hole is shown in Figure 18.Figure 17:A cross section of the G3 target plate through both impact locations​1138Figure 18:Image of the center hole (Shot #2) on the G3 sample. The blue part in the image is a​fragment of the sabot.The G4 system was also investigated. Here, the G4 has the same thickness of Kevlar as G3, but all the reinforcement is contained in the middle of the plate. Post-impact images of the target plate can be seen in Figure 19. The figure shows that the projectile produced a large exit damage zone on this plate, indicating that this plate may be near being able to stop the projectile.

The figure shows that the projectile produced a large exit damage zone on this plate, indicating that this plate may be near being able to stop the projectile. The cross section of the G4 specimen is shown in Figure 20 where plastic deformation of the Kevlar is clearly seen. The figure also shows delamination at the interface between the Kevlar and the Onyx, an effective mechanism for dissipating energy. This is a thicker delamination than seen in the other architectures. A higher magnification image of the penetration location is shown in Figure 2, where it is observed an intralaminar delamination within the Kevlar fiber. Here, the hole is narrow and no spall cone were displayed.Further optical analysis was carried out on the impacted zone from the top view with a Keyence microscope using 2D and 3D capabilities (see figure 21).

Figure 21 shows a delamination bulge at the entry hole, a feature attributed to a shockwave deflection from Kevlar layer.1139Figure 19: Impacted G4 target plate. Front face (left), and back face (right).​Figure 20:A cross section of the impacted G4 target plate.​Figure 21:A microscope image of the penetration zone in the G4 target plate. The blue​fragments are from the sabot.1140Figure 21:A microscope image of the penetration zone in the G4 target plate from the top view​(left). The blue fragments are from the sabot. 3D rendering at right.4. ConclusionsThe following conclusions can be made from the Charpy impact testing:The Z orientation is noticeably weaker than the XY Notch Down orientation regardless ofOnyx infill pattern type examined.The XY Notch Side orientation did not see a significant difference in impact strength between layer thickness of 0.100 mm and 0.200 mm.

For XY Notch Down, the impact strength of 0.200 mm is greater than that of 0.100 mm and 0.125 mm. The difference in [14] may be the result may be the result of the Onyx findings from previous research being more brittle than nylon, a performance associated with the presence of the chopped fiber reinforcement within Onyx.Kevlar reinforcement provides no useful role if the impact notch is parallel to the Kevlarreinforcement planes as in the case in the reinforced Z orientation specimens.In the XY Notch Down orientation, the Kevlar reinforcement layer is perpendicular to the notch plane then more energy is required to fracture the specimen. (see architecture #16 with a layer of 1.49 cu.cm of Kevlar versus architecture #15, the impact energy leaped from 0.832 lb ft to 8.394 lb ft!). The fracture path becomes exceptionally tortuous as the crack will attempt to progress by delaminating the reinforcement layer away from the Onyx and tearing fiber when possible.

The fracture path becomes exceptionally tortuous as the crack will attempt to progress by delaminating the reinforcement layer away from the Onyx and tearing fiber when possible. This is similar to observations from previous research involving nylon (not Onyx) reinforced by Kevlar .In the case of the XY Notch Side orientation, the addition of Kevlar is of some value but not dramatic. In the XY Notch Side orientation, only the edge of the reinforcement layer is presented to the crack and tearing and bypassing of the Kevlar is easier at this ​In the case of the XY Notch Side orientation, the addition of Kevlar is of some value but not dramatic. In the XY Notch Side orientation, only the edge of the reinforcement layer is presented to the crack and tearing and bypassing of the Kevlar is easier at this ​1141orientation.

In the XY Notch Side orientation, only the edge of the reinforcement layer is presented to the crack and tearing and bypassing of the Kevlar is easier at this ​1141orientation. This also agrees with observations from previous research involving nylon reinforced by Kevlar[14].​​While the highest performance was achieved by Kevlar reinforced specimens with rectangular Onyx infill, both cost and machine time are high for such a combination. A desirable mix of cost, weight, and performance was obtained by HEX Onyx infill patterns with combined ISO and CON reinforcement fill patterns).The following conclusions can be made from the gas gun projectile impact testing:From the difference in velocity testing,the exit damage zone will be small when overmatch is high. The exit damage zone will be large when overmatch is small which means the architecture is close to stopping the projectile at that target thickness and projectile velocity.

The exit damage zone will be large when overmatch is small which means the architecture is close to stopping the projectile at that target thickness and projectile velocity. Based on this observation, G1 and G2 architectures were highly overmatched. G3 and G4 architectures were less overmatched. Increasing the plate thickness to 8 mm or 10 mm and while keeping the relative % percentage and location of reinforcement should result in the projectile being stopped at approximately 255 m/s. ● For a sparse infill, the entry hole of the round will be less than the diameter of the impact projectile as the structure will elastically expand and then contract. However, the exit zone will be larger due to plastic deformation and spall. There will be an exit damage zone where the surface layer on the back side of the target will delaminate and even tear off.Architecture is important.

There will be an exit damage zone where the surface layer on the back side of the target will delaminate and even tear off.Architecture is important. If the Kevlar is positioned near the surface of the front or impact side of the target and there is no reinforcement near the exit or back side of the target, there will be a spall cone. If there is reinforcement near the back side of the target, the spall cone is eliminated.Surface delamination appears to follow raster patterns and form an X shape. The depth of delamination can be a function of the depth of the Kevlar layer. In architecture G3 the Onyx delamination layer extended 0.5 mm deep to the Kevlar layer. Architecture G4 had a thicker delamination layer that extended to the Kevlar layer which was 2 mm deep.Another general conclusion is that Charpy V-notch testing is an imperfect predictor of gas gun performance. The Charpy geometry has a notch that extends across the width of the specimen. The gas gun impact is localized.

The Charpy geometry has a notch that extends across the width of the specimen. The gas gun impact is localized. What works well in Charpy might not work well in gas gun impact. For Charpy specimens, reinforcement needs to be close to the notch in order to increase impact energy. For gas gun targets, the reinforcement should be away from the impact site in order to prevent the formation of a spall cone. Regardless of method, a key finding is that architecture determines the performance against impact.1142References1. Yang B. Blade containment evaluation of civil aircraft engines. Chin J Aeronaut. 2013;26: 9–16.2. Kim H, DeFrancisci G, Chen ZM, Rhymer J. Impact damage formation on composite aircraft structures. UCSD FAA JAMS Paper, Technical Review Meeting. 2012. Available: https://www.niar.wichita.edu/niarworkshops/Portals/0/UCSD_Kim_PP.pdf3. Richardson WJ, West T. Serious birdstrike accidents to military aircraft: updated list and summary.

plementation and Evaluation of Novel Buildstyles in FusedDeposition Modeling (FDM)Tobias Wasser, Anshu Dhar Jayal, Christoph PistorUniversity of Utah, Salt Lake City, 19991 Abstract Previous investigations have shown that the optimization of extrusion dynamics in .con-junction with the buildstyle pattern is of paramount importance to increase part quality inFused Deposition Modeling (FDM). Recently domain decomposition and space fillingcurves have been introduced for slice generation in FDM [1]. The current work focuseson the implementations of fractal-like buildstyle .patterns using. Simulated Annealing [2,3], Lin-Kernighan algorithms [4] and Construction Procedures based on NearestNeighbor Heuristics [5]. These computational optimization procedures are able togenerate filling patterns that allow the continuous deposition of a single road to fill arbitrary shaped domains.

These computational optimization procedures are able togenerate filling patterns that allow the continuous deposition of a single road to fill arbitrary shaped domains. The necessary software modules to produce arbitrary threedimensional artifacts have been developed and are evaluated with respect to part qualityand build time.2 Introduction Today path· generation and optimization problems found in manufacturing processes,such as FDM, typically aim for solutions that results in the shortest possible cycle time.On one hand this necessitates the determination of the shortest path. On the other hand,the manufacturing cycle time will be minimized the higher the speed which cah·beachieved on the individual path segment~. As a result, variations in traveling speed areincreased whenever.

As a result, variations in traveling speed areincreased whenever. sharp turns are taken, due to the deceleration and accelerationlimitations of<motion control and••·the manufacturing>equipment.iAllmanufacl11ringprocesses are more or less sensitive to these changes in traveling speed but usuallyrequire more advanced control capabilities as· speed variations increase.·· Additionalexamples of.processe~ particularly sensitive·to tooltravelingspeedvariations·wouldbesurface. Both also ·benefit from a non-self-intersecting continuouspath,,,,hich is a mustfor the FDM process. The simplest solution to the problem is therefore to apply/hatch95patterns in Meander-form, which consist of parallel lines only, that fillthe desireddomain. However, this approach results in a highly oriented final structure, which canonly be by changing the Meander-orientation in different slices.

However, this approach results in a highly oriented final structure, which canonly be by changing the Meander-orientation in different slices. Some of the problemscaused by this orientation are:IIhighly anisotropic parts (Figure 1),IIdiscontinuous nozzle- or toolpath,IIvoids due to minimum curve radii and areas which cannot be reached byMeander-like patterns andIIpatterns of the surface structure/roughness (surface- or microstructure).A solution to this problem is a path with segments of random orientation, which cancompletely cover the domain to be processed and which does not self-intersect. SpaceFilling Curves (SFC) such as the Hilbert SFC have been used in the past [1] to build partswith the FDM process. But until now these processes were limited to highly regularshaped objects and not applicable for arbitrary boundaries.20 , - - - - - . - - - - - , - - - - - - , - - - - , . . - - - - - r - - - - , - - - - , - - - - - - ,181614m a..

- - - - - , - - - - - - , - - - - , . . - - - - - r - - - - , - - - - , - - - - - - ,181614m a.. ~ ~ iii1210..!!l 'iii ai I-8656Strain [%]Figure 1: Stress-Strain Diagram for Uni-Directional (En' E 22,E33) and Nearest Neighbor Heuristics (NNH) Tool-Path96With the introduction of underlying equidistant grids that are able to fill arbitraryboundaries, it is possible to get a number of nodes in each plane, which must be visitedby the nozzle or tool. The spacing of the grid for the FDM-process is based on theroadwidth of the deposited material. For the Stratasys FDM 1650 a minimum roadwidthof 0.508 mmhas been measured. The schedule of the visit must fulfill certain criteria,which are similar to the definitions for Eulerian and Hamiltonian tours in graph theory[5]:1. all nodes have to be visited,2. no node must be visited more then once,3.

all nodes have to be visited,2. no node must be visited more then once,3. minimal distance between nodes,4.shortest possible total path length (sum over distances),5.there may be no intersecting path elements,6. only vertical and horizontal moves and7. only in the nodes at the boundary diagonal moves are allowed.The computational solutions for the problems with the aforementioned requirements arebased on the Traveling Salesman Problem (TSP) [3,5]. The general TSP is not limited topoints on equidistant grids but allows weight factors to discribe the "distances" betweennodes or better criterias. Therefore the TSP is also adressing problems beyond Euclideanplanes or even problems were the distance between AB isnot the same as the distancefrom BA (i.e. the distance matrix is not symetric).3 Solutions for the Traveling Salesman ProblemState of the art solutions to the Traveling Salesman Problem are based on solutions tooptimization problems.

the distance matrix is not symetric).3 Solutions for the Traveling Salesman ProblemState of the art solutions to the Traveling Salesman Problem are based on solutions tooptimization problems. After a first random tourthat visits all nodes only once andallows intersections, the total length (sum of distances) is optimized, using the fact that ashorter tour· always exists when crossings. are still present [5]. Very simple optimizationalgorithms are based on node or edge· insertion algorithms. Better quality algorithms arebased for example on Simulated Annealing or theLin-Kernighan algorithm. Instead ofusing random tours at the beginning of optimization algorithms it is also possible to useheuristics to develop •start up solutions.

Instead ofusing random tours at the beginning of optimization algorithms it is also possible to useheuristics to develop •start up solutions. The described Maze .algorithm is based onNearest Neighbor Heuristics and makes use of Rotation Operations.973.1· Simple Optimization Algorithms Node insertion algorithms select nodes whose position in the tour schedule is changed.TAGi!lPde§<tP~q~a~gl}daJl~t~\)ir.~~0~p§iNPn are chos\)n complete1yxandOmly and unguided. Edge insertion. functionsar;~si11lilar to the simple ·node in$ertiou functionexcept, thattwo nodes get selected. Then the order of alLpoints in between gets reversed.And the new edge is inserted in front of another selected node of the remaining tour. Forsmall problems, this random procedure is able. to achieve optimum solutions. But it alsocan happen that the algorithm becomes trapped in local minima.

Forsmall problems, this random procedure is able. to achieve optimum solutions. But it alsocan happen that the algorithm becomes trapped in local minima. With an increase inproblem size (number of nodes)the improvementin unguided random selectiondecreases asymptotically. Therefore all of the better algorithms focus on either guidingthe solution finding process or on solving the problem of local minima.3.2 Simulated Annealing The algorithm starts at a high-energy (cost or initial tour) configuration. Then thesimulation chooses random neighboring configurations called Careful Annealing inphysical systems and Simulated Annealing in optimization problems. Ifthe neighboringconfigurations result in an improvement in the Energy State (tour length),then thesimulation is updated, in order to achieve the Ground·State or the Optimal Solution [2,3].Until now this procedure matches the simple optimization algorithm.

However byintroducing a probability factor which judges if a configuration is defined as the new stateeven if this neighbor is not better then his origin configuration, the algorithm solves theproblem of being trapped in local minima. This probability factor is also continuouslyupgraded during iterations [2,3].Problematic during the use of Simllltaneous Annealing is the number of nodes; evenwith long running times there is the possibility that the Optimal Solutions intersect, sinceintersections are not optimized. LaarhovenandAa,rts [2]· recommended to run the wholealgorithm as often as there are nodes, in order .to secure an .optimal tour safely. Thisincreases the. running. time by the number of nodes.

LaarhovenandAa,rts [2]· recommended to run the wholealgorithm as often as there are nodes, in order .to secure an .optimal tour safely. Thisincreases the. running. time by the number of nodes. When the initial configuration isconstructed by nearest/neighbor heuristics, it is noticed that the annealing < process firstworsens the configuration and cannot.guarantee an optimal solution.3.3 Lin-Kernighan Algorithm Lin and Kernighan introduced their improvement heuristic in 1971 [4]. Their obj.ectivewas to improve the toolpath of a numerically controlled laser used for cutting holes. Inorder to improve the cycle time, their focus was on the positioning of the laser and not onthe cutting process, since the laser was only burning holes with equal diameters. WhileSimulated Annealing focused on the problems of trapped situations, Lin and Kernighanstarted with guiding the acquisition of neighboring solutions.

WhileSimulated Annealing focused on the problems of trapped situations, Lin and Kernighanstarted with guiding the acquisition of neighboring solutions. The main idea was thatsince the single distances caused in their sum the total cost or length of eachconfiguration, ~n optimization algorithm should find better positions fornodes with verylong single distances. ThereforeLin and Kernighan used guidingcriteria in orderto reach an6optimal solution. Throughout the5literaturetheLin-Kernighanalgorithm is recognized as theheuristic with the most promisingresults.In orderto solve the2problem of getting trapped inlocal minima and to achievebetterrunning times, Lin andKernighanfoundthebestsolutions by using five nodesol=================:-J o 10 6 5 Number of Iteration Figure·2: Total.patblel1gthoverthe numbers of iter ation for ~problem withl0S9711od.es{tensile..bar).234789x 104simultaneously in each improvement loop.

Similar algorithms are called k-optimizationalgorithms whereby k is thenumber of nodes in each loop. The improvement rate or time gain over all other processesis noticed throughout all problem instances, but similar to all discussed heuristics thissolution rate is highly asymptotic (Figure 2).3.4 Heuristics for initial configurations Instead of optimizing a.complete random initial configuration, heuristics can be used. tofind better startup solutions. Both Insertion Heuristics and Nearest Neighbor Heuristics99have been implemented and evaluated.

tofind better startup solutions. Both Insertion Heuristics and Nearest Neighbor Heuristics99have been implemented and evaluated. Since the implementation of Nearest NeighborHeuristics fulfill the requirements of non-self intersecting pathelements, an algorithmusing rotation operations has been designed to create complete toolpaths.3.4.1 Insertion HeuristicsStarting with an initial randomly chosen tour containing only a small amount of nodes,the final path is obtained by adding new nodes to this tour according to certain criteria.The starting tour consists of three nodes,which thus had as a Hamiltonian tour theminimal possible total path length. The newnodes were selected either randomly or bythe criteria of maximal(Figure 3) orminimal distance to the rest of the alreadyconfigured nodes. After their selection thenodes were included in the tour such thattheir position was a minimal addition in thetotal path length.

After their selection thenodes were included in the tour such thattheir position was a minimal addition in thetotal path length. These heuristics cancreate very good initial configurations inshort time, since the running time is O(n),Figure 3: Tour consisting of 325 nodes achieved by Farthest Insertion Heuristic.and they can create or introduce tour solutions that are similar to meanders (thereare longquasi-parallel sections); however intersections still occur.3.4.2 Nearest Neighbor HepristicsIn these algorithms a·· single random node is selected and from this starting node, theentire tour is constructed by adding one new node at a time. The objective for each newnode is that it has.a minimal distance to the current end pointof the tour. For the equidistant grid this means that there aretheoretically four possible connections But there are only upto three new possibilities, since the last connection reachingthis point cannot be opened or lost.

The main problem is thatthis algorithm gets trapped in areas, where all of theFigure 4: A dead end.surrounding nodes are already part of the tour configuration.100In other words the last node of the tour has reached a dead end (Figure 4).The solution for this dead end is to allow the algorithm to connect randomly or guidedto a free node. Butthe results are not acceptable asintersections occur in the generated toolpath. However, ifthis algorithm is to be used for finding a first initial tour thatwill be optimized by another algorithm, the selection of freenodes in this fashion is tolerable. In fact the quality of thetour configuration is already so good that the subsequentlyFigure 5: Rotationimplemented ()ptimization algorithms always lead to longer Operationtotal path lengths.Instead of "jumping" in the tour to another free node, it is also possible to rotate theconnections.

The Operations that changed the configuration shown in Figure 4 to theconfiguration in Figure 5 were:1.selecting a node which is already part of the tour configuration,2. opening its connection with the next node in the tour schedule,3.turning the scheduled order of points around for those nodes, which werefreed from the tour schedule and4. connecting the trapped node (circle) from before to the selected node.With these changes a new tour end is achieved and the Nearest Neighbor Heuristic startsagain. An example in Figure 6 showsthe broad variation in a solution for a1,300 nodes problem. A simple distancealgorithm guides the selection of nodeswhere the connections are being brokenup to the nearest free nodes. And in thefinal version,thealgorithm allowsdiagonal moves at the boundary.

And in thefinal version,thealgorithm allowsdiagonal moves at the boundary. On theStratasys FDM 1650, severaltensilebars were built, in which the toolpathwas configured by the Nearest NeighborFigure 6:.1,3QOnOde toura.chieved through Nearest Neighbor Hem:istic.Heuristic. One slice of a standard tensile bar includes 10,597 nodes. The subsequently101conducted Stress-Strain tests revealed ahigher Elastic modulus (Figure 7). Thebuilt parts had a higher density. and theirfracture behavior was close to injectionmolted ABS parts.

M. Pistor, S. I. GUceri, "Domain Decomposition and Space FillingCurves in Toolpath Planning and Generation," University of Illinois at Chicago, 1998.(2)P. J. M. van Laarhoven, E. H. L. Aarts, "Simulated Annealing: Theory and Applications," D.ReidelPublisching Company: Dordrecht, 1987.[3)Kunwoo Lee, "Principles of CAD/CAM/CAE Systems," Addison-Wesley: Reading, Seoul, 1999.[4)S. Lin, B. W. Kernighan, "An Effective Heuristic Algorithm for the Traveling-Salesman Problem,"Bell Telephone Laboratories: Murray Hill, 1971.[5)Gerhard Reinelt, "The Traveling Salesman: Computational Solutions for TSP Applications,"Springer-Verlag: Berlin, Heidelberg, 1994.102.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperImplementation of 3d printer in the hands-on material processing course: An Educational PaperB.(cid:3)Silwal*Mechanical Engineering Department, Georgia Southern University, Statesboro, GA, 30458AbstractManufacturing equipment and technology will continue to be an integral part of engineering education. Using advanced and modern equipment and technology not only emphasizes hands-on activity-based teaching and learning, but it also makes the educator and the program competitive. There are many advantages to implementing 3-d printers in engineering education; however, there are also some challenges.

There are many advantages to implementing 3-d printers in engineering education; however, there are also some challenges. In this paper, a two-year(cid:3) implementation of 3-d printers in a traditional material processing undergraduate engineering course has been(cid:3) presented and discussed.Introduction******************************************************************************************Additive manufacturing a.k.a. 3d printing is a buzz word in the manufacturing workplace. There are many advantages of the 3d printer; however, the most significant advantage is the ability to fabricate a more complex part that has never been realized via another manufacturing process. Thus, it brings design flexibility to engineers and students who may not need new knowledge of the manufacturing processes that are usually required to design a complex part.

Thus, it brings design flexibility to engineers and students who may not need new knowledge of the manufacturing processes that are usually required to design a complex part. With the evolution of materials and application-specific 3d printers, it has created a paradigm shift for educators and engineers as the term ‘design for manufacturing’ is being replaced by ‘manufacturing by design.’Manufacturing technology and equipment has been introduced in many engineering labs’ curriculum to improve education and provide opportunities for hands-on learning. Modern engineering technologies are an essential tool to add in engineering education to make the program and the educators competitive. Implementing 3d printing in engineering education is highly beneficial to the students as well as the teacher (can also use instructor instead of teacher). It has been used as a rapid prototyping tool in the past.

It has been used as a rapid prototyping tool in the past. Stereolithography based printers were widely available in universities in the early 21st century; however, it was not feasible to implement towards the educational activities, particularly in teaching labs. With the invention of fused deposition modeling (FDM) based3-d printers and reduction in the cost of 3-d printers, it is now affordable to implement 3-d printers in educational activities, mainly in teaching labs. There are many pre-engineering and engineering curricula for 3d printing skills that have been developed to teach in high schools and universities [2, 3, 7]. A detail information on where and how the 3-d printers and additive manufacture are implemented in schools and universities are well documented in [8].

A detail information on where and how the 3-d printers and additive manufacture are implemented in schools and universities are well documented in [8]. Although the core implementation is still in rapid prototyping of a new design and outreach activities, it has also been supplemented in several engineering courses in our university, such as:• machine design to compute gear-ratios [5] • mechanism design to visualize a mechanism • components of a senior design projectengineering graphics course to improve the student’s spatial skills45Sand casting is one of the most used and essential casting processes in the world. A cavity enclosed inside the sand mold determines the shape of the part. The cavity is formed by packing the sand around a pattern. The pattern is a replica of the object to be cast and is generally made out of plastic, metal or wood. The casting metal is heated slightly above its melting temperature.

The pattern is a replica of the object to be cast and is generally made out of plastic, metal or wood. The casting metal is heated slightly above its melting temperature. The liquid metal is then poured into the cavity in an open mold casting. The metal is poured through the down sprue and via the gating system (runner) in a closed mold casting as depicted in Figure 1.The quality of casting depends upon many attributes: the thickness of the pattern, pouring temperature, pouring rate, gating system, riser, etc. The pattern is usually a replica made out of wood, metal, or plastic, while the riser compensates the shrinkage that generally occurs during the casting process [1]. Foundrymen (people who work in a foundry) design these patterns based on experience and practice. 3-d printing technology has been implemented in the hands-on material processing course at Georgia Southern University. The 3-d printed parts are used as patterns in the casting industry for the casting of metal parts.

The 3-d printed parts are used as patterns in the casting industry for the casting of metal parts. The 3-d printed wax pattern has the potential to replace the expensive and time-consuming current wax pattern used in the casting process. Advances in technology have allowed the 3-d printing of sand molds [4] which would not even require a pattern. Ford has used the 3-d printed sand molds for casting prototype engines, cylinder heads, etc. [6].

Ford has used the 3-d printed sand molds for casting prototype engines, cylinder heads, etc. [6]. The implementation of the 3-d printer for pattern making process in the undergraduate level course will bring creativity in design, improve the students’ critical thinking ability and teaches the fundamentals of metal casting.l PouringRiserCast metalDownsprueCast metal(_ _ _J~==k=~==t----l=-=-::,..r--~:---::---7 Parting lineRunnerL_ ____ _CoreOpen MoldClose MoldFigure 1 : An open mold and close mold sand casting processCourse description ****************************************************************************************** MENG 3333 (course name and number) is a junior level ‘material processing’ 15 weeks semester-long course for mechanical engineering students. The course is 2-2-3 level course, which means 2 hours lecture and 2 hours lab are assigned in a week, and students earn three credits.

The course is 2-2-3 level course, which means 2 hours lecture and 2 hours lab are assigned in a week, and students earn three credits. The junior level ‘material processing’ course teaches the students traditional and modern material processing techniques. The main objective of the lab is to teach the students hands-on-work on key manufacturing processes. The course objectives are: 1. Demonstrate workplace safety and material and tools handling safety 2. Teach how to create a bill of materials, plan of procedures, working drawings for machine shop production 3. Teach the major applications of a variety of material processing techniques, and how these processes are used in manufacturing4. Recognize the fundamentals of major manufacturing processes and the process physics465. Teach how to estimate the cost of manufacturing processes 6.

Recognize the fundamentals of major manufacturing processes and the process physics465. Teach how to estimate the cost of manufacturing processes 6. Demonstrate the process and use examples and videos to reinforce the key principles in material processing7.Demonstrate an understanding of the contents and use of machining reference and handbooks and other appropriate material resourcesThe student course outcomes are:1. Identify, define, and categorize equipment, tooling, and operations used in the processing of engineering materials.2. Compare and contrast materials processing techniques and delineate the limitations of each as they relate to manufacturing.3. Describe the effects of processing on the properties of materials. 4. Explain the influence of materials and processing parameters on manufacturing processes and operations. 5. Conduct analysis of forces, pressures, friction, wear necessary for the processing of a variety of materials.

5. Conduct analysis of forces, pressures, friction, wear necessary for the processing of a variety of materials. Communicate information pertaining to design considerations, product quality, and cost factors for each type or group of manufacturing processes. 6. processing techniques and processing equipment used to manufacture goods from these materials Discuss the four types of engineering materials: metals, ceramics, polymers, and composites and theLab project ****************************************************************************************** MENG 3333, a lecture and lab integrated project-based course, covers the applications and use of different discipline areas of manufacturing processes, such as metal-casting processes and equipment, forming and shaping processes and equipment, joining methods and equipment, molding, extrusion and fabrication of polymers, and composite processing techniques.

The blade and fasteners are acquired through a vendor. The frame and retainer are made using sheet metal bending. The tensioner is made using CNC and manual machining. The students will learn sand casting, manual and CNC machining, sheet metal bending, resistive welding, grinding, etc in the lab and lecture settings. From the lesson learned, students are required to design and fabricate a hacksaw handle during the semester. The knowledge gained from the previous Solid modeling course (mold design) enables them to design the solid pattern in CAD format. The students then 3d print their design. The main criteria are to leverage the advantage of the 3d printer to group-specific handle design and reduce the volume without compromising the processing and structural integrity. Also, design flexibility can be leveraged out of the 3d printer.

Also, design flexibility can be leveraged out of the 3d printer. A detailed lab schedule is presented in Table 1.------------------------------■ Table 1: A semester-long schedule of the material processing labWeek 1Introduction to lab safety, equipment, and toolsWeek 2Demonstration of the Sand casting processWeek 3Design of Hacksaw handleWeek 4Design of Hacksaw handle/3-d printingWeek 5Assembly and engineering drawingWeek 6CNC – G Code IntroductionWeek 7CNC machining/Lathe machiningWeek 8CastingWeek 9CastingWeek 10CastingWeek 11CastingWeek 12Shaping, Bending, Welding, etc.Week 13Shaping, Bending, Welding, etc.Week 14Post processingWeek 15Presentation and Project submissionEvery semester, 100-120 students enroll in the junior level material processing course. Each lab covers 16-20 students.

Each lab covers 16-20 students. There has been a significant investment made by the college to acquire different brands and different scale of 3d printers for research and teaching purposes. Figure 3 presents the varieties of 3d printers at the college additive manufacturing lab.One of the Accreditation Board for Engineering and Technology (ABET) new student outcomes is that the engineering program must demonstrate that the students can attain: ‘an ability to apply engineering design to produce solutions that meet specified needs with consideration of public health, safety, and welfare, as well as global, cultural, social, environmental, and economic factors.’ The student's project in MENG 3333 is ideal for assessing such an outcome. The students design the hacksaw using the principles of ergonomics, low weight, defect-free higher strength handle.

The students design the hacksaw using the principles of ergonomics, low weight, defect-free higher strength handle. An alternative process Silicone mold casting (cold casting or resin casting) is also used to compare with the sand casting process ass(cid:72)(cid:86)(cid:86)ing the environmental and economic factors. Silicone48mold casting is a resin casting process where silicone is poured over the additively manufactured pattern to create a silicon mold. A resin is then poured inside the mold to create the desired part.The students are divided into four groups, with each group containing four students to foster the teamwork. Each group conduct the market survey and design a computer-aided design (CAD) model of the hacksaw handle in Solidworks. A complete sand mold making process using an old pattern (not 3d printed) will be demonstrated before the design process. The instructor fully laid out the requirements to have quality casting.

A complete sand mold making process using an old pattern (not 3d printed) will be demonstrated before the design process. The instructor fully laid out the requirements to have quality casting. The main reason for this is to integrate metacognition and critical thinking skills in the design process. As the pattern design is dependent upon the mold making and casting process, the demonstration will allow the student to think beyond the box and come up with a better design.Figure 3: The 3d printer lab, Uprint models on the left, Fortus 450 printer in the middle, Color resin printer on the right and small scale rep-rap 3d printers in the middleDesign for casting******************************************************************************************Draft angle, liquid, and solid shrinkage, etc. have to be considered while designing the handle. The pattern will have two halves.

have to be considered while designing the handle. The pattern will have two halves. The students are required to understand the sand-mold making process entirely to design the two halves of the pattern considering the draft angle, and shrinkages,49Shrinkage: The actual size of the cavity is slightly oversized to allow for the shrinkage. Due to the thermal expansion coefficient, most metal and alloys shrink during cooling.Draft angleRiser: The riser (Figure 1) acts as a reservoir and solidifies after the main casting, thus negating the shrinkage in the casting. To compensate the liquid shrinkage in the sand casting process and if the casting is very thick, a riser is design and implemented.Draft: A draft angle (1°) is required in the pattern (shown in Figure 4) to eject the pattern from the sand effortlessly. A draft angle assists or aids in the removal of the pattern from the sand.

A draft angle assists or aids in the removal of the pattern from the sand. A thin film exists between the sand mold and the pattern which creates a vacuum, thus damaging the sand mold while removing the pattern from the sand. A negative mold also tears the mold apart during the pattern removal process.Figure 4 : The draft angleSplit Pattern: Two-piece pattern is the most widely used pattern for sand casting. One half of the pattern is molded in drag and the other half in cope.Sizing: The 3d printer is capable of printing anything less than 8” x 8”.Creativity in Design and Design Freedom ***************************************************************************************** * Some of the students’ pattern designs are presented in Figure 6. A clearance fit between the two split-patterns is required in the design to fit two-patterns together during the mold making process.

A clearance fit between the two split-patterns is required in the design to fit two-patterns together during the mold making process. The top-right pattern (Figure 6) is a simple design of a split pattern where a male end meets with the female end. Critical thinking plays an important role in design, however no critical thinking was not observed in this design; first the weight is not controlled, second no draft-angle has been introduced, and it breaks the sand when pulling out from the sand mold. If the sand gets pulled from the mold, it generates undesirable defects on the casting handle.The top-left design (Figure 6) is very similar; however, some weight has been removed to reduce material consumption and time in printing. The bottom-left split pattern (Figure 6) is more creative as both the pattern contains the boss, however, the boss is so thin that it broke just after printing.

The bottom-left split pattern (Figure 6) is more creative as both the pattern contains the boss, however, the boss is so thin that it broke just after printing. The design of a boss is essential in sand casting as it won’t break any sand during the mold removal process as depicted in Figure 5. The pattern also has excess material removed, thus it prints quicker than the other patterns. The bottom right pattern (Figure 6) resembles the grip for holding the handle, design based on the principle of ergonomics.Figure 7 shows a pattern with a core in place for the sand casting. As seen in the Figure, it is thicker than the other pattern presented in Figure 6 and generates defects in the casting and requires multiple risers to cast the handle without defects. Ideally, it is easier to cast with silicone mold resin casting than sand casting. There is a trade-off between sand casting and silicone casting process. First, the silicone casting process is less expensive than sand casting.

There is a trade-off between sand casting and silicone casting process. First, the silicone casting process is less expensive than sand casting. Second, the silicone casting process requires minimum effort. Third, the silicone resin casting process doesn’t require heating and melting; thus, it may have some environmental benefits. The sand casting, on the other hand, brings aesthetic appeal to the parts being aluminum and has higher strength. The sand casting is the most used process so the fundamental knowledge of heating, cooling, solidification, etc.

The sand casting is the most used process so the fundamental knowledge of heating, cooling, solidification, etc. is ideal for introducing in a junior level course.50Figure 5: Representation of the boss in the pattern to easily removed the pattern from the sand moldFigure 6 6: Several split patterns design ed by students51Figure 7 7: A design that requires a core for sand casting and multiple risers, however easier to resin cast without any y r e problem******************************************************************************************Procedure/Investment/Cost ****************************************************************************************** A price sheet form was designed and placed to implement the 3d printer across the different departments inside the college. The students fill out the form with specific information about the volume and weight of the pattern.

The students fill out the form with specific information about the volume and weight of the pattern. The form along with the STL file is emailed to the 3d printer lab where a student worker will perform the printing operation.

The standard material price is $2.45 per in3 ( per 16.38 cm3) and $4.76 for the support material. These are calculated using the material and depreciation cost of the 3-d printer. Considering 10-20 % of the part volume is used as support material, the total cost of printing the split pattern is approximately $75.Challenges in implementing 3d printer in educational settings *****************************************************************************************Some of the new factors that were never introduced before the implementation of the 3d printer were the draft angle and solidification shrinkage. An old pattern made out of aluminum was used in every lab. Other processing techniques, such as machining and injection modeling, were quite complex and time- consuming to fabricate a pattern.

An old pattern made out of aluminum was used in every lab. Other processing techniques, such as machining and injection modeling, were quite complex and time- consuming to fabricate a pattern. It was a very cost-effective solution at that time.Figure 8: The surface topography comparison between 3-d printer and cast partsThe 3-d printed parts are not within the specified tolerance range. The printer has ±0.001” tolerance; however, a tolerance of 0.002 will not provide enough clearance between the hole and peg of the52split pattern. The support material surface finish is not smooth enough for casting. Figure 8 provides a comparison between the 3-d printed pattern and sand cast hacksaw handle. As seen in Figure 8 the surface topography is retained in sand casting. One of the essential factors to implement 3-d printing in engineering education is time and reliability.

As seen in Figure 8 the surface topography is retained in sand casting. One of the essential factors to implement 3-d printing in engineering education is time and reliability. If the students are involved in project-based learning with a fixed timeframe, the reliability of the 3d printers is an important issue. The general requirement is to have minimum maintenance with the 3d printers, however not all 3d printers are consistent with this remark. We have more success with U-print 3d printers without much need for maintenance within the warranty period. Now the warranty period has phased out, they are starting to show their age, and we're starting to have more significant problems. The periodic maintenance and service of parts such as: belts, hot ends, and material drive bay etc. is required. The parent company wants $2800 per year, per machine for the warranty service. Considering total of six printers the cost is relatively high.

is required. The parent company wants $2800 per year, per machine for the warranty service. Considering total of six printers the cost is relatively high. Most of the educational facility (particularly the university) has budgeted for equipment acquisition; however, there is not sufficient operational budget for the maintenance of out-of-service high-cost 3d printers. Figure 9 shows the out-of-service printer. It is obsolete, and we can't buy parts for it.An alternative to sand casting, cold casting or resin casting is a process where a silicone mold is used to make castings from the resin. The surface finish of such casting is far better than the print out of FDM based 3d printer. Figure 10 shows the silicone mold and resin cast hacksaw handle. This process has also been used to compare with the sand casting process in terms of environmental, economic, and safety factors.

Figure 10 shows the silicone mold and resin cast hacksaw handle. This process has also been used to compare with the sand casting process in terms of environmental, economic, and safety factors. The students compare the sand casting with the silicon mold casting process in terms of cost, safety, environment factor etc. and make a presentation on their findings.Figure 9 : A out of service 3dd e 3 d printer ready to recycledFigure 10: A silicone mold to use for resin casting processsStudents Response ********************************************(cid:13)(cid:13)******************************************** 53A survey was designed to collect the students’ response (n = 29) on 3d printer use. Based on the results, nearly 50% of the students haven’t used a 3d printer or 3d printer related service before this course (Figure 10). 6.9% owned a 3d printer before taking this course. 68.97 % rated the quality of the 3d printed product reasonable.

6.9% owned a 3d printer before taking this course. 68.97 % rated the quality of the 3d printed product reasonable. 75.86 % appreciated the 3d printing integration in this course, while the rest presumed that it is common to have 3d printer activities in engineering course. In terms of 3d printing services, 20.69 % find it challenging to access 3d printing services.

In terms of 3d printing services, 20.69 % find it challenging to access 3d printing services. 100 % responded that they would like to be involved in 3d printing activities postgraduation given an opportunity (Figure 10).YesYesNoNo0%10%20%30%40%50%60%70%80%90% 100%0%10%20%30%40%50%60%70%80%90% 100%Figure 11 : Students re sponse on “Have you used 3 - d printer or services prior to this course?” on left and “ “Will you involve in 3 - d printing activities post t- graduation given an opportunity? ” on right tConcluding Remarks: ******************************************************************************************In this paper, the implementation of 3d printers in a material processing course has been presented. The use of 3d printers has helped us to teach both traditional and modern manufacturing processes. Some of the factors that were never introduced before the implementation of the 3d printer were introduced in pattern design.

Some of the factors that were never introduced before the implementation of the 3d printer were introduced in pattern design. With the implementation, the students have a better practical knowledge of pattern design for sand casting and this paper proves that the engineering design is influenced by manufacturing technique. The students’ response suggests they are enthusiastic about the use of 3d printers and its integration in the curriculum. High cost 3d printers need a high operational budget in an educational facility to self-sustain after the warranty period. It is of best interest to acquire intermediate-cost, low maintenance 3d printers rather than high-cost high end 3d printers for teaching purposes.54References1. Groover, M. P. (2007). Fundamentals of modern manufacturing: materials processes, and systems. John Wiley & Sons.2. Chien, Y. H. (2017). Developing a pre-engineering curriculum for 3D printing skills for high school technology education.

PLEMENTATION OF A FUNCTIONALLY GRADIENT MATERIAL MODELING AND DESIGN SYSTEMAparajit Pratap, Richard H. Crawford Laboratory for Freeform Fabrication, Department of Mechanical Engineering The University of Texas at Austin, Austin, TX 78712 Reviewed, accepted August 13, 2003 AbstractNew advancements in Solid Freeform Fabrication (SFF) processes promise the capabilityto produce Functionally Gradient Material (FGM) parts, in which the material compositions vary spatially. To realize this potential there is a need for CAD methods and design software to model, design, represent and exchange material information and instructions to the manufacturing process. However, currently available commercial CAD systems are limited to representing and storing only geometric information, which is not adequate for material design purposes.

However, currently available commercial CAD systems are limited to representing and storing only geometric information, which is not adequate for material design purposes. This work presents an extension of a theoretical approach based on Volumetric Multi- Texturing (VMT) and hypertexturing schemes to make the material design process intuitive and user controllable. Inverse distance weighted interpolation is used in conjunction with procedural material functions to accomplish axial or linear material gradient directions from surface to surface across a solid. This offers the capability of specifying fixed material composition values to the faces in the solid and blending them across the interior of the solid. The extension of the proposed approach to the modeling of discrete material domains is also discussed. These material regions can be combined using special sets of operators depending on the form of the material functions.

These material regions can be combined using special sets of operators depending on the form of the material functions. Finally, a design environment has been developed, which allows users to systematically apply material information to geometry and captures design intent.1 IntroductionSolid Freeform Fabrication encompasses emerging rapid prototyping technologies usedin making components with complex geometries without human intervention or part-specific tooling. SFF processes involve the use of a computer-generated model to provide the geometric information required in the fabrication process. In general SFF is based upon layered manufacturing techniques, in which the parts are produced layer by layer in an additive fashion. SFF thus has access to the interiors of solids and provides the most suitable means of building parts with highly accurate local material composition control, which cannot be achieved by conventional fabrication processes.

These capabilities enable the use of the technology in the manufacture of parts composed of functionally gradient materials. Functionally gradient material components have smoothly varying material composition. They possess large variations in the volume fractions of their constituents and variations in microstructure because of which they exhibit some useful properties such as reduced thermal and residual stresses at joints, improved thermal and mechanical properties and high specific strength and oxidation properties [4].The current state of solid and geometric modeling is at an advanced stage of developmenthowever existing commercially available CAD modelers do not support material design information. Thus there is a need for new modeling systems that can design and represent heterogeneous material parts. The problem addressed in this work is the implementation of a150design environment for multi-material part modeling and representation.

The problem addressed in this work is the implementation of a150design environment for multi-material part modeling and representation. The methodologies used are largely based on implicit procedural and Volumetric Multi-Texturing (VMT) schemes [9]. Issues such as combining and blending the effects due to multiple material constituents at a point, extensibility of smoothly varying material gradients to modeling of discrete material interfaces and achieving multiple directionality of material variation within a part have been addressed. Appropriate data structures for geometry and material attributes have been constructed for material data storage and retrieval and user interface design issues have been discussed, with emphasis on user interaction and input methods to make the design process more intuitive and predictable for users.1.1 Previous WorkSeveral researchers have been working to develop heterogeneous material modelingsystems.

Kumar and Dutta [7] proposed a set-based approach for spatial discretization of the solid interior by associating material density functions with material subsets called rm sets. Pegna and Safi [12], Kumar and Wood [6] and König and Fadel [5] have used finite element based methods to model and optimize material density distributions given a particular design objective and constraints. Marsan and Dutta [8] and Jackson and Petrikalakis [3] have applied volumetric mesh schemes to describe material data.Several mesh free methods have also been applied to FGM design, which do not rely onany form of spatial decomposition of the geometry. Hu and Jiang [2], Siu and Tan [15] and Wahlborg and Ganter [16] are some researchers who have adopted heterogeneous solid modeling techniques to material modeling.

Hu and Jiang [2], Siu and Tan [15] and Wahlborg and Ganter [16] are some researchers who have adopted heterogeneous solid modeling techniques to material modeling. Wahlborg and Ganter [16] have introduced a heterogeneous implicit solid modeling scheme that makes use of the Boolean operators union, intersection and difference on heterogeneous objects to model both solid as well as material spaces defined as implicit primitives. Biswas and Shapiro [1] discuss the parameterization of the interior of a shape using functions of normalized distance fields from material features with known material properties. The material functions are continuous and constructed using known design and manufacturing constraints and are interpolated using transfinite interpolation, with inverse distance weights to obtain overall material functions.Representations in the form of spatial discretizations are inefficient due to large data sizesand errors due to approximations.

This adds to the cost of computation and to the difficulty in modifying material models. The schemes implemented in this work are based on a mesh free implicit procedural approach and are therefore devoid of the errors that arise from spatial decomposition.2 Implicit Procedural ApproachThe material modeling scheme used in the implementation of the design interface islargely based on implicit procedural and Volumetric Multi-Texturing techniques that were first introduced by Park et al [9].1512.1 HypertextureHypertexture is a technique in computer graphics that makes use of 3D volume density functions to create effects such as explosions, furry objects and eroded surfaces or objects that have variations in density. The functions that control the density distribution in the fuzzy region are called density modulation functions (DMF) and are chosen to be higher order procedural functions built upon bias and gain functions as described by Perlin [13].

Using the VMT approach any pointr intp 3G can be interrogated to return the material densityipd at that point using a setgeometric space of procedural functions mF such that for each material i: ,r ( i pF mr p)=∈3i pdG,intintn∑r ( i pF mintn) ∑ =di p=1i=1i=1The densitiespd physically represent volume fractions of each material and sum to one functions, mF ’s are1−nat each point for all n materials that compose the part. Therefore only needed to fully constrain the model. Each function, mF is a procedure consisting of several that are attached to surfaces (faces or group of faces) in the part contributing functions,imfgeometry.

Each function, mF is a procedure consisting of several that are attached to surfaces (faces or group of faces) in the part contributing functions,imfgeometry. These contributing functions are combined together using the blending function described in equation (5) and are functions of the distances of the sample point from the part surfaces and reference geometry such that the closer the sample point is to an entity the greater is the influence of the contributing function associated with that entity in determining the density at that point.Park [11] introduces two classes of gradient functions, namely, global and local (surface)gradients. Global gradients are one-dimensional scalar functions (either density modulation functions or user-defined functions) defined over material spaces called blobs that are mapped onto 3D geometry. The shape of the global material space determines the overall pattern of the material distribution.

The shape of the global material space determines the overall pattern of the material distribution. In this work a B-rep representation of the material space geometry has been added to the existing implicit surface description by Park [10]. If the material function is a function of an independent parameter t, in the case of B-reps, t varies linearly with distance from the boundary surfaces whereas in the case of an implicit surface, t evaluates to the implicit function value and is therefore proportional to the distance from the surface. Material gradient functions similar to those used with global gradients when assigned to the surfaces of the part geometry serve as surface gradients.

Material gradient functions similar to those used with global gradients when assigned to the surfaces of the part geometry serve as surface gradients. The global gradient function, when multiplied with the surface gradient, is trimmed to force the net gradient to conform to the boundary surface of the part geometry.3 Multi-Directional Material VariationThe existing material modeling method focuses on improving surface coatings on components that are exposed to severe thermal and mechanical stress environments. The current approach therefore places importance on material variation from the surfaces of components into the interior and does not support general material distributions in the interior. Further, user-153(6)prescribed boundary material compositions on surface features are not satisfied.

Further, user-153(6)prescribed boundary material compositions on surface features are not satisfied. Modeling a linear material gradient pattern that starts with one material fraction on one face and blends into another material fraction on another can be achieved by using multiple surface gradient functions for each face, one material function for each material. Consider an object with different material functions assigned to each face in a pair of adjacent faces. A simple blend of the two functions in the overlapping region where each face has an influence on the material properties at points on the adjacent face, would not yield the exact material compositions specified on the faces.

For this reason the functions for each material for each face need to be trimmed to the boundary values on the neighboring face using a different approach, as described below.3.1 Inverse Distance Weighting FunctionConsider n faces or surface features and let the material composition of a material k be≤≤iC on the face iF , such that point p due to the influence of all n faces is:0iC1. The net material composition of material k at a givenfk(p)=fblend, k(p) .n∑i= 1( pWC ii)⎟wherefblend, •kiW is a weight function controlling the influence of the blended material function, )p ( In order forwith the face iF.

These functions are called the inverse distance weighting functions or Shepard’s method of data interpolation, as discussed by Biswas and Shapiro [1].iu is the distance of p from the faceiF and the exponent k controls the smoothness of the154(7)(8)(9)(10)The limitation of this approach is that it does not ensure that the sum of the materialdensities at a point p for all materialsm∑( j pf)equals 1, where m is the total number of materials.=1One way to treat this inadequacy is to consider the remaining quantity1m− ∑j= 1( j pf)toconstitute porosity in the part or to constitute the volume fraction of the ( Another way is to normalize the volume fraction due to an individual material k to obtain: ()thm 1+( ) pf k ( )∑ pf j)=fpNk ,mmaterial.j=1such thatm∑fNk ,(p) 1 =k=1As a result of this addition, it is now possible to model multiple material gradient patterns, namely linear as well as concentric or radial material gradations so as to achieve any direction of material variation.

Figure 1 shows concentric contours of material gradation between two materials for a cross-section of a bottle mold. It is modeled by assigning a single density function for the primary material to each of the four faces forming the boundary, which blends into the secondary material in the interior. In Figure 2, three faces have been assigned 100% composition of one material each. Using the proposed method, it is now possible to clamp each material composition to be exactly 100% on its respective face and achieve a gradual blend of the three into each other. Here the depth of material influence for each face is set to be greater or equal to the length of a side of the square cross-section.Figure 1: Concentric Pattern Figure 2: Linear gradation patternwith 3 materials4 Modeling Discrete Material DomainsThe current approach does not account for discrete heterogeneity of materials and henceoffers only a restricted case of material modeling.

Modeling discrete material regions can be considered as a special case of FGM modeling. The same approach can be applied to discrete material modeling by using constant functions or step functions with the reference geometry. The global material spaces can be used to define discrete material domains confined within the B-rep by assigning constant material fractions to individual blobs. Since assigning a value of 1 to the155(11)(12)bias and gain functions makes the density modulation function in equation 3 a step function at the midpoint of the [0, 1] interval, material compositions could be made to change discretely at these points (see Figure 3). The blobs can then be combined using regularized Boolean operations for B-rep geometry and min and max functions [16] or equivalent Ricci’s functions for implicit solids.

Consider the drill insert shown in Figure 4. The two spheres illustrate the use of blobs as global density gradients to define discrete material regions. Figure 5 is obtained after performing a difference of the upper sphere from the bottom one. The net gradient takes the shape expected if a Boolean subtraction were performed on the spheres.156(13)The smooth gradient near the boundary is the result of using smooth functions for the surface gradient.5 Implementation and ResultsThe material modeler has been developed and implemented using the 3D ACIS® Modelerand the HOOPS 3D Application Framework. ACIS is a geometric modeling kernel written in C++ that supports modeling of wireframe, surface and solid geometry in a unified data structure. It defines geometric shapes as B-reps and saves model information in the “SAT” format, a format that is slowly becoming a de facto standard for the exchange of CAD data.

It defines geometric shapes as B-reps and saves model information in the “SAT” format, a format that is slowly becoming a de facto standard for the exchange of CAD data. ACIS has been used to build the data structures for the geometry and material information and to perform the required modeling operations. A skeletal application based on HOOPS 3D, a graphics toolkit, has been used to design the user interface for the application.5.1 Attribute Architecture for Material InformationACIS supports user-defined attributes that can be attached to geometric and topologicalentities and can be saved along with them in an SAT file. In this research, user defined attribute classes constructed in ACIS have been used to store the material data associated with the part for use by procedures that generate material distributions. Figure 6 shows the data structure used for modeling the overall density gradient in a part.

Figure 6 shows the data structure used for modeling the overall density gradient in a part. The blocks in blue represent the user defined attribute classes and the arrows, their associations with each other.The blob attribute (ATTRIB_BLB) consists of either an ACIS BODY (B-rep) or a userdefined implicit function describing an implicit solid primitive since the global density gradient can be modeled in either one of these two ways. This attribute also consists of the bias and gain parameters for the density modulation function or a user defined material function associated with the blob surface. Finally it specifies the depth of influence of the blob surface over the material composition at an interior point.The blend attribute (ATTRIB_BND) located one level higher in the hierarchy storesinformation on how to combine the blob functions using union, intersect or difference operators as described in section 4.

It also contains a blending factor, a, which is used to blend the blob functions in the case of a union operation using equation 5. The blend attribute is assigned to the part to be modeled. The blob attributes are stacked together in a doubly linked list such that at all times the blob at the top of the stack is pointed to by the owning entity, which is the blend attribute in this case. The remaining blobs can be accessed using the “next” pointer of the previous one in the list. The blend attribute in turn is owned by the ACIS BODY and can be accessed by it.The surface gradient data is stored in two separate attribute classes, map and join. The map attribute (ATTRIB_MAP) contains an array of material compositions assigned to a face on the part and arrays of dmf parameters of bias and gain for each material function attached to the face. The owning entity in this case is an ACIS FACE (see Figure 7).

The owning entity in this case is an ACIS FACE (see Figure 7). There is also a provision for user defined material functions for each material and an array for the depths of influence of the face for the contributing material functions for each material.157ACIS B-REP (BODY) (Owner for BLEND)BLEND (ATTRIB_BND) (Owner for BLOBs)NULLBLOB 1 (ATTRIB_BLB)BLOB 2BLOB NFigure 6: Data Structure for Global Density GradientOne of the features of the material modeler implementation is the automatic selection offaces to be used to blend associated material functions in the part. The approach used here is that of blending faces that are adjacent to an edge or vertex, using a blending factor unique to that edge or vertex.

The approach used here is that of blending faces that are adjacent to an edge or vertex, using a blending factor unique to that edge or vertex. For this purpose a join attribute class (ATTRIB_JON) containing the value for the blending factor is created and attached to an EDGE or VERTEX owning entity (see Figure 7).FACE (Owner)EDGE (Owner)VERTEX (Owner)MAP (ATTRIB_MAP)JOIN (ATTRIB_JON)Figure 7: Data Structures for Surface Density GradientsThe user interface (GUI) for the material modeling application has been built using the HOOPS 3D Application Framework (HOOPS/3dAF). It supports the implementation of the material modeling features that have been described above. B-rep models in the ACIS SAT format can be imported in the viewer. The interface has the provision for creating solid primitives and performing Boolean operations on them to serve as complex geometry for the blobs.

The interface has the provision for creating solid primitives and performing Boolean operations on them to serve as complex geometry for the blobs. Implicit functions for the definition of blob geometries and material density functions for both global and surface gradients can then be assigned to the part geometry. Finally geometry and material attributes can be saved and retrieved to and from SAT files. The computed material distribution is displayed as an image.158NULL5.2 ResultsThe following example is the FGM design of a pulley shown in Figure 8. Here the design intent is to make the part more wear resistant in areas subject to maximum wear. For example the pulley can be built from stainless steel while having a greater concentration of a wear resistant material like carbide in areas near the hub and the rim [3].The part is modeled using a surface gradient approach.

The material information for onlyone material (primary material) is entered (carbide in this case) and the secondary material density (stainless steel) is the complement of the first. The faces that compose the surfaces of the rim and hub of the pulley respectively, are selected and assigned a material function (see Figure 8 (a) and 8 (b)). Since these surfaces are composed of the same material (carbide) they are all assigned a 100% material fraction for the same material. Figure 8 (c) displays the resulting material gradient distribution on a cross-section of the pulley passing through its center laterally.(a)(b)(c)Figure 8: FGM Design of a pulleyThis example illustrates the ease with which the model of the pulley can be designed for FGM variation. The design is shown to be completed in four steps using the material modeling interface.

The design is shown to be completed in four steps using the material modeling interface. It can also be seen that the material variation conforms to the original design intent of having greater concentration of carbide at the hub and the rim, indicated by the region colored in magenta and a gradually increasing concentration of stainless steel with increasing distance away from the hub and the rim on the spokes.6 Conclusions and Future WorkNew developments in SFF processes promise the capability of fabricating functionallygradient material parts. To realize this potential new CAD methods and design software must be developed to enable designers to model, design, and transmit FGM objects and instructions to the manufacturing process. FGM’s are expected to show clear advantages over homogeneous materials in terms of mechanical and thermal properties.

FGM’s are expected to show clear advantages over homogeneous materials in terms of mechanical and thermal properties. The goals of the work presented were to take existing research [9] based on implicit procedural methods and extend it in order to build a tool to design volumetric material information accurately and intuitively.

The advantages of the procedural approach are accurate description of three-dimensional material data and compactness in data storage by having to store only the instructions for the procedures and their arguments.159Volume discretization methods are discarded, leading to increased accuracy, concise representation and resolution independence.Inverse distance weighted functions have been used in conjunction with the proceduralmaterial functions to accomplish axial or linear material gradient directions and offers the user the capability of specifying fixed material composition values to the faces in the solid and blending them across the interior of the solid. The definition of global density gradients can be expanded by representing them as both implicit surfaces as well as boundary representations (B- reps). This permits the user to draw or import more complex geometry to define global material spaces.

This permits the user to draw or import more complex geometry to define global material spaces. However the use of B-reps to define global gradients significantly increases the amount of data that must be stored compared to a single compact implicit function expression that could define an entire implicit solid. It has been proposed that one-dimensional density functions can be step functions or constant material fractions in order to represent homogeneous material regions. These material regions can be combined using special sets of operators depending on the form of the material functions. Finally, a usable material modeling interface has been designed by which designers can systematically apply material information to B-rep geometry and observe the resulting material distributions in cross-sections of the solid.

The user interface captures design intent and is an intuitive tool for designers to use.6.1 Future WorkApart from material volume fractions in a component, the macro-structures are equally significant in determining the physical properties of the part. Therefore designing and optimizing the structure variation of materials, e.g. composites, is a valid area of further research. This therefore presents the possibility of designing and optimizing smooth spatial (3D) variations in a structure between two materials by grading reinforcement orientations. Another prospective area for exploration is the formulation of a design methodology for FGM objects that must be placed in the context of an overall design process. Creating such a design methodology would involve investigating how a designer interacts with the model. For FGM modeling, the specification of graded material information is closely related to the geometric design.

For FGM modeling, the specification of graded material information is closely related to the geometric design. Ideally, a design system would allow the user to interactively modify the geometry and the composition in any order. Methods for achieving this goal depend on the underlying representation as well as the chosen design methods.Though geometric data can be exchanged to a limited extent between data formats thereis still a need for a new standard data exchange format that is universally supported by most CAD systems to efficiently and accurately translate non-geometric data and attributes from one format to another. In addition to this there would also be the need for new data translators to translate data back and forth between native formats and the new standard format. The next step in multi-material manufacturing is to use the material model and convert it into machine instructions by developing an SFF process-specific process-planning algorithm.

The next step in multi-material manufacturing is to use the material model and convert it into machine instructions by developing an SFF process-specific process-planning algorithm. The research can only be validated when results in the form of physical products can be produced. Therefore the development of the necessary hardware and the compatible software for tool-path generation is a vital research area that needs to be investigated in the future.160References1. Biswas, A., Shapiro, V., Tsukanov, I., “Heterogeneous Material Modeling with Distance Fields”, Technical Report SAL 2002-4, Spatial Automation Laboratory, University of Wisconsin-Madison, 1513 University Avenue, Madison, WI 53706, June 2002.2. Hu, X., Jiang, T., Lin, F., Sun, W., “Reasoning Boolean Operation for Modeling, Simulation and Fabrication of Heterogeneous Objects”, Proceedings of the 2000 Solid Freeform Fabrication Symposium, Austin, TX, August, 2000, pp. 417-427.3. Jackson, T. R., Patrikalakis, N.

34, Issue 10, September 2002, pp. 705-716.16. Wahlborg, J., Ganter M. A., Schwartz, D. T., Storti, D., “H-ISM: An Implementation of Implicit Solid Modeling”, ASME Design Engineering Technical Heterogeneous Conference, Montreal, Canada, September, 2002.161.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperImplementation of tophat profile laserinto low temperature process of Poly Phenylene Sulfide Fumio ITO, ASPECT Inc.Toshiki NIINO, Institute of Industrial Science, the University of TokyoAbstractPoly Phenylene Sulfide is a high performance plastics that is used in high-value parts. Themost commercially available plastic laser sintering machines employ Gaussian profile CO2 laser.In this research, low temperature process of the material with tophat profile laser is tested. Byusing tophat profile, a tensile strength of 64.5 MPa was obtained. The value is equivalent to 75%of those from injection molding and 140% of standard laser sintered polyamide.

Byusing tophat profile, a tensile strength of 64.5 MPa was obtained. The value is equivalent to 75%of those from injection molding and 140% of standard laser sintered polyamide. Tophat profileeffectively suppressed sparkling and fuming and improved surface finish decreasing surplus sintering or sticking powderIntroductionIn recent years, application of plastic laser sintering (LS) to end-parts production is expandingin aerospace and medical industries. Although these applications often require high performanceplastics possessing high strength, high heat resistance and/or biological compatibility, process ofthese plastics has not been commercialized, or is extremely expensive. This limitation in materialperformance restricts application range of LS.One of the problems with LS process is part warpage caused by layer-by-layer shrinkage.

This limitation in materialperformance restricts application range of LS.One of the problems with LS process is part warpage caused by layer-by-layer shrinkage. Toavoid this, in commercially available systems, powder bed is preheated so as to keep itstemperature in a range known as “process window,” which is between recrystallization and meltingtemperature [1]. Although this powder preheating effectively prevents warpage, it causes anotherproblem. Generally, plastic does not melt or recrystallize at a certain point of temperature but in arange. When the ranges are close to each other, the temperature control is quite difficult, or thewindow even closes. Additionally, preheating requires LS machine with a high-heat-resistancewhen we use high performance plastics due to their high melting temperature. Poly Ether EtherKetone (PEEK) is one of the high performance plastics that has the highest performances.

Poly Ether EtherKetone (PEEK) is one of the high performance plastics that has the highest performances. Inprocess of PEEK, powder bed temperature must be kept above 300°C [2] and this preheatingelevates the machine price and deteriorates the powder recyclability.Previously, the authors introduced and tested a novel LS process which prevents part warpageby fixing a part to a rigid base plate instead of preheating powder bed at such a high temperature[3, 4]. In the following, we call this novel process and standard one as “low temperature process”and “high temperature process.” Additionally, the authors implemented this technology forprocessing of PEEK, and succeeded in prototyping several parts at a low bed temperature of 200°C2194[5].Poly Phenylene Sulfide (PPS) is also a high performance plastics. Though PPS is slightly inferiorto PEEK in terms of physical performance, its price is much cheaper than PEEK.

Though PPS is slightly inferiorto PEEK in terms of physical performance, its price is much cheaper than PEEK. As a result, thereare many products made of PPS [6], and this shows great demand for PPS as a material of high-value functional parts. Technically speaking, we can expect that LS process of PPS is easier thanPEEK because of its lower melting temperature.In this research, low temperature process of PPS is tested. Sparkling and fuming, which is alsofound in process of PEEK, is suppressed by reducing peak intensity of light. For this purposetophat profile laser is employed. Quality of obtained parts in terms of mechanical strength andprecision is discussed. Although tophat profile laser is utilized in selective laser melting of metalpowder [7], no plastic LS machine using this laser is commercialized yet.

Although tophat profile laser is utilized in selective laser melting of metalpowder [7], no plastic LS machine using this laser is commercialized yet. Advantage of usingtophat profile laser in plastic LS is also discussed.Material and MethodMaterialPPS powder having median diameter of 75µm was employed. Median diameter was measuredwith a particle size analyzer (MT3300, MicrotrackBEL Corp.). Thermal and mechanicalperformance of an injection molded part of non-reinforced PPS was summarized in Tbl.1 [8].Optical micrograph of PPS powder grains is shown in Fig.1Tbl.1 Performance of material (typical example of PPS)Type of materialTrue densityMelting pointPPS(Poly Phenylene Sulfide) 1.35 g/cm3 275 °CTensile strength85 MPaFig.1 Optical micrograph of PPS powder grains2195Laser sintering apparatus and tophat profile laserTwo experimental apparatuses are employed. CO2 lasers are installed to both machines.

CO2 lasers are installed to both machines. Oneis equipped with Gaussian profile (or single mode) optics, and the other with tophat profile. Thebeam spot diameter of Gaussian profile is adjustable from 130µm to 340µm by changing thedistance between focusing lens and powder bed. The beam spot diameter of the tophat profile is480µm. To compare the influence of profiles under the same laser power, beam spot diameter ofGaussian profile laser is adjusted to 228µm so that the two profile share the same intensity gradientat the beam edge as shown in Fig.4. A pair of galvanometers is able to scan the beam in x and ydirection in a range of approximately 100×100mm. To inhibit oxidation of fabricated parts, oxygenconcentration of building chamber is lowered below 0.1% by nitrogen gas purge. It was unable tomake parts of PPS with high temperature process because of the highest powder bed temperature of these apparatuses is 200°C.

It was unable tomake parts of PPS with high temperature process because of the highest powder bed temperature of these apparatuses is 200°C. An injection molded PPS plate is used as the base plate.Fig.2 Laser sintering system2196Focusing lens (Z-Galvanometer)CO2 laserX-Y GalvanometerIR thermometerZnSe laser windowFeed heaterRecoatingrollerPart bed heaterOver flow pocketPart cylinderFeed pistonFeed cylinderPart pistonFig.3 Schematic view of the laser sintering system200 µm200 µmFig.4 Profiles of Gaussian and tophat profile laserVisual observation and measurementTo evaluate quality of fabricated parts, density and tensile strength were measured. Density isused as an index of strength. Fracture surface of parts were visually observed to help understandingof sintering mechanism. Sparkling and fuming during process are the main obstacle of process.

Fracture surface of parts were visually observed to help understandingof sintering mechanism. Sparkling and fuming during process are the main obstacle of process. Tosuppress them, surface of the bed was visually observed to know the cause of these phenomena.To evaluate process precision, side surface of parts is visually observed.These observation of fracture and side surface for evaluations was carried out with SEM(VHX-2000 and VHX-D510, KEYENCE). Tensile test pieces were fabricated and measure withuniversal testing machine (Series 5560, INSTRON). Size of test pieces was shown in Fig.5 andconformed to ISO 527-2. Tensile speed was 2mm/min. Density measurement was carried out byArchimedes principle using electronic balance (AX120, SHIMADZU Corp.) for weight2197measurements.805823050 1Fig.5 Size of tensile test pieceBuild parameter of each laserPowder bed temperature was maintained 200°C and layer thickness was set to 100µm inprocess.

Build parameters of Gaussian and tophat beam were shown in Tbl.2. Scanning speed andinterval were adjusted so that the two profiles give the same energy per unit area.Tbl.2 Build parameter for the Gaussian and tophat beamProfileGaussianTophatLaserpower11 W11 WScanningspeed0.92 m/s1.0 m/sScanninginterval100 µm90 µmEnergy perunit area 119 kJ/m2 122 kJ/m2Beamdiameter228 µm480 µmExperimental ResultsFabrication process and density with Gaussian profile laserFig.6 shows powder bed in fabrication process of PPS with Gaussian profile laser. Fig.7 showssparkling and fuming in fabrication process of PEEK[5]. In fig.6, fuming was not observed,however sparkling was observed which was not extremely big.

Fig.7 showssparkling and fuming in fabrication process of PEEK[5]. In fig.6, fuming was not observed,however sparkling was observed which was not extremely big. Relative density was obtained 92.6% (1.25g/cm3).SparkleFig.6 Sparkle in process with Gaussian profile laser2198Fig.7 Sparkling and fuming in process of PEEK[5]Fabrication process and density with tophat profile laserFig.8 shows powder bed in fabrication process of PPS with tophat profile laser. Fuming wasnot observed either. Sparkling was suppressed successfully in comparison with Gaussian profilelaser. Small sparks were observed yet, but it is not as much as it can obstacle the process or lower the part quality. Relative density was improved to 95.6% (1.29g/cm3.)SparkleFig.8 Sparkle in process with tophat profile laserTensile strengthResults of tensile strength measurement were summarized Tbl.3 and Fig.9. Tensile strengthtests were carried out three time per each profile laser.

Tensile strengthtests were carried out three time per each profile laser. Average tensile strength of test piecesfabricated with Gaussian and tophat beam were 47.2MPa and 64.5MPa, respectively. Using tophatprofile improved tensile strength by a factor of 1.37, although almost the same amount of energy2199was supplied per unit area.Tbl.3 Tensile strength of parts fabricatedProfileGaussianTophatEnergy perunit area 119 kJ/m2 122 kJ/m2Tensilestrength47.2 MPa64.5 MPa70] a P M[ h t g n e r t s60504030e l i s n e T20100GaussianTop hatFig.9 Results of tensile strength testFracture surface and side surfaceSEM images of fracture surfaces are shown in Fig.10, 11. On the fracture surface of partsobtained by Gaussian profile, many linear horizontal cracks are observed. The cracks occur at aninterval of 100 or 200µm.

On the fracture surface of partsobtained by Gaussian profile, many linear horizontal cracks are observed. The cracks occur at aninterval of 100 or 200µm. Although we can see cracks on the fracture surface of parts from tophatprofile, they are much shorter, not horizontal and nonperiodic. SEM images of exterior sidesurfaces of parts are shown in Fig.12, 13. The part obtained by tophat profile is smoother on theside surface. Amount of powder grains sticking on the surface looks smaller when tophat is used.2200Fig.10 Fracture surface of the part obtained with Gaussian profile laserFig.11Fracture surface of the part obtained with tophat profile laserFig.12 Side surface of the part obtained with Gaussian profile laser2201Fig.13Side surface of the part obtained with tophat profile laserDiscussionIn previous research, sparkling and fuming during processing PEEK were suppressed byexpanding beam diameter to reduce intensity of light without decreasing laser power.

In thisresearch, it was found that tophat profile can suppressed the phenomena better. The reason for thisimprovement is that tophat profile reduces the peak intensity and supplies constant energy over itsbeam spot while Gaussian profile is strong in the middle and weak at the outer edge. Tophat profilemakes improvement in terms of mechanical strength as well. Although the supplied energy perunit area is the same, tensile strength obtained by tophat profile is 1.4 times of what Gaussian beamcan provide.Tensile strength of parts obtained by tophat profile in this research was 64.5MPa. This isequivalent to 75% of injection molded parts and 140% of standard laser sintered polyamide (PA)[4, 5].

This isequivalent to 75% of injection molded parts and 140% of standard laser sintered polyamide (PA)[4, 5]. Considering that PPS has a better anti-heat performance than PA12 or 11, we can concludethat low temperature process of PPS can make dramatic progress in plastic laser sintering.By the both of Gaussian and tophat profile laser, a high relative density more than 92% wasobtained. The parts obtained by Gaussian profile includes long, horizontal and periodic cracks.The periodicity of the cracks and their interval, which is equivalent to layer thickness of 100µm orits integer multiple, shows that these cracks derive from boundary of layers. In contrast, partsobtained by tophat profile does not include trace of layers. This indicates tophat profile can providebetter interlayer connection.

In contrast, partsobtained by tophat profile does not include trace of layers. This indicates tophat profile can providebetter interlayer connection. It is surmised that light of tophat profile laser or heat generated by thelaser penetrates into deeper place without causing other phenomena like sparkling.Between tophat and Gaussian profile, we found differences in roughness and an amount ofsticking powder of the exterior surface. Cause of these differences is not clear. Throughinvestigation into definition or precision should be done in the near future.2202ConclusionPPS powder was successfully processed with low temperature process. It was confirmed thatusing tophat beam is effective to implement the process and obtain parts having high tensilestrength by suppressing violent phenomena such as sparkling and fuming. Gaussian and tophatprofile laser provide the same density when amount of energy per unit area is the same.

Gaussian and tophatprofile laser provide the same density when amount of energy per unit area is the same. However,tensile strength of parts obtained by tophat profile laser was 1.4 times as great as by Gaussianbeam.Tophat profile is effective to suppress sparkling and fuming. The profile also provide the partswith smooth exterior surface decreasing surplus sintering or sticking powder.

plementing Rapid Prototyping Using CNC Machining (CNC-RP) Through a CAD/CAM InterfaceMatthew C. Frank, PhD Department of Industrial and Manufacturing Systems Engineering The Iowa State University of Science and Technology Ames, Iowa 50010 mfrank@iastate.eduAbstract:This paper presents the methodology and implementation of a rapid machining system using a CAD/CAM interface. Rapid Prototyping using CNC Machining (CNC-RP) is a method that has been developed which enables automatic generation of process plans for a machined component. The challenge with CNC-RP is not the technical problems of material removal, but with all of the required setup, fixture and toolpath planning, which has previously required a skilled machinist.

The challenge with CNC-RP is not the technical problems of material removal, but with all of the required setup, fixture and toolpath planning, which has previously required a skilled machinist. Through the use of advanced geometric algorithms, we have implemented an interface with a CAD/CAM system that allows true automatic NC code generation directly from a CAD model with no human interaction; a capability necessary for a practical rapid prototyping system.1. IntroductionMost commercial RP systems are based on additive processes whereby models are constructed by stacking 21/2-D cross sectional layers on top of one another. The additive RP systems are often limited in both geometric accuracy and material quality. Subtractive processes such as CNC machining have advantages over the limited choice of materials and the limited functionality of parts produced by additive processes.

Subtractive processes such as CNC machining have advantages over the limited choice of materials and the limited functionality of parts produced by additive processes. However, machining is not a completely automated method in either the process or fixture planning steps. There has been a need for a rapid machining system, but previous attempts to automate CNC machining have been approached from the perspective of traditional machining methods. It has become necessary to re-think how parts can be held, oriented, and then actually cut; perhaps borrowing methods from existing approaches to additive rapid prototyping processes.Traditional machining requires extensive planning by a specialized and experienced machining technician. Moreover, the challenge of machining complex and intricately shaped components is daunting even on the most advanced machines.

Moreover, the challenge of machining complex and intricately shaped components is daunting even on the most advanced machines. In traditional machining, the focus is typically on simpler geometries (holes, slots, planes, etc), or, when the geometry is more complex, as in an airplane wing, the shape is defined by known geometric functions. In these cases, a skilled machinist chooses cutting tools based on the desired shape (such as a drill bit for a round hole). With respect to this research effort, Rapid Prototyping and Manufacturing presents an enormous challenge in terms of geometric complexity and every part build is unique.In manufacturing, the time spent developing an optimal process plan for machining a component can be easily absorbed over a long production run. However, if truly one-of-a-kind components are required, then the engineering involved in planning and implementing the solution must be automated or eliminated altogether.

However, if truly one-of-a-kind components are required, then the engineering involved in planning and implementing the solution must be automated or eliminated altogether. Process planning for CNC machiningincludes tasks such as determining: 1) how to clamp the stock material (fixture planning), 2) what orientation to approach the geometric surfaces with the cutter (setup planning), 3) what size and shape cutter to use (tool selection) and finally, 4) how to traverse the geometry with the cutting tool in order to create the surfaces (toolpath planning) [M1,C1,J1]. Computer Aided Process Planning (CAPP) systems have made some reduction in the time required to plan machining operations, but the time, skill and cost for one-of-a-kind machined parts is still dominated by the planning before machining and not the machining itself. There are currently no approaches that are flexible enough to provide a universal solution for such various shapes as would be needed in rapid prototyping.

There are currently no approaches that are flexible enough to provide a universal solution for such various shapes as would be needed in rapid prototyping. To this end, CNC-RP is a new rapid machining process that has been developed to allow for push-button machining of complex parts with no advanced manufacturing skill required and no process planning time after the CAD model is obtained.The CNC-RP process has been realized as a module in the MasterCAM® package that automates virtually all of the process and setup planning tasks. Using several geometric algorithms and standardized tool and material libraries, this system can process a CAD model similar to the way an RP software interface processes an STL file. The interface computes setups, creates sacrificial support structures, generates toolpaths and outputs the required NC code and setup instructions.

The interface computes setups, creates sacrificial support structures, generates toolpaths and outputs the required NC code and setup instructions. To process a part, the user simply loads the prescribed diameter and length stock material into the CNC machine and downloads the NC code for processing. Hence, CNC-RP has been implemented as one of the first truly push-button subtractive rapid prototyping systems, able to be implemented on a wide array of CNC milling machines.2. CNC-RP MethodologyThis rapid machining process is based on a setup strategy whereby a rotary device is used to orient round stock material that is fixed between two opposing chucks. Rotating the stock using an indexer eliminates the inherent problem of retaining reference coordinates associated with re- clamping a part in a conventional fixture. For each orientation, all visible surfaces are machined and a set of sacrificial supports keep it connected to the uncut ends of the stock material.

For each orientation, all visible surfaces are machined and a set of sacrificial supports keep it connected to the uncut ends of the stock material. Once all operations are complete, the supports are severed (sawed or milled) in a final series of operations and the part is removed. The setup and steps to this process are illustrated in Figure 1. As an example in this figure, a component is being machined using sacrificial supports to retain the part at its ends along the axis of rotation. This method of using one axis of rotation for indexing between setups is obviously not capable of machining all parts of extremely complex shape. Parts with severely undercut features or complex features on three or more mutually orthogonal faces may not be machinable with this approach.

Parts with severely undercut features or complex features on three or more mutually orthogonal faces may not be machinable with this approach. In particular, this setup strategy assumes that some axis of rotation exists such that all surfaces are visible.The rapid machining process incrementally creates the part by machining layer by layer for each orientation, thus it is not very fast and would not be a good choice for production manufacturing. However, this process is extremely rapid because it can almost automatically begin processing a part directly from a CAD file.

However, this process is extremely rapid because it can almost automatically begin processing a part directly from a CAD file. The research so far has developed algorithms that automatically analyze the part geometry and determine 1) if the part is machinable, 2) if so, how many orientations are required and 3) all the process parameters including the stock size, tool size, feeds, speeds, sacrificial support geometry and the sequence of operations This new rapid machining process is novel in that it approaches the planning steps in a “feature-free” mannerusing robust automated methods. That is, geometric algorithms calculate the steps necessary to machine all the surfaces of the object.

Implementing CNC-RP in CAD/CAMThe CNC-RP process involves several sequential steps that need to be automated in the CAD/CAM environment. This paper will briefly introduce each step, but the details of the algorithms will be neglected due to space considerations, but can be found in several publications [F1,F2,L1,L2]. The focus of the paper will be to illustrate how each geometric and/or process planning algorithm was implemented in the CAD/CAM environment.The process is initiated after the CAD model of the part is loaded in the CAD/CAM system. Next, a sequence of steps is initiated including 1) analyzing the part for an axis of rotation 2) establishing a work coordinate system 3) attaching sacrificial supports 4) determining setup orientations about the axis, 4) generating a roughing process to remove the bulk of material and 5) generating a finishing process to create the surface geometry.

Borrowing from traditional RP, this system does use the STL file format; however, not for creating the final NC code. In this research, we use the STL file for a feature-free analysis of the CAD model that enables us to determine all the process and fixture plans. Several geometric algorithms and heuristic tools have been developed to work with either the STL file or a slice file from the STL. However, the original native CAD file is preserved through the process. As such, the CAD model provides the drive surfaces for toolpath planning, which eliminates the tpical error associate with the STL format. Depending on the step in the process, we export varying “quality” STL files, by changing the chordal deviation of the output. This is done because we can greatly reduce the computational load on the geometric algorithms, when precision is less important. As shown in Figure X, there one path flow for the CAD model and then several “offline” analysis steps that are performed on STL files.

As shown in Figure X, there one path flow for the CAD model and then several “offline” analysis steps that are performed on STL files. The flowchart’s upper chain illustrates the steps in the CAD/CAM environment, while the lower chains represent the offline algorithms. steps of the process and implementation are presented below.Imported CAD modelOriented CAD modelSupports attached to CAD modelCAD model rotated and toolpaths createdNC CODESTL modelSTL modelSTL modelSETUP SHEETAxis analysisSacrificial support designSetup orientation analysisTool and parameter selectionUser inputUser inputUser inputFigure 2 – System Flowchart illustrating interaction between CAD models in the CAM system and algorithms, through the STL file formatThe following sections below will describe the individual steps in this system.

Eachsection roughly corresponds to a step in the process of turning a CAD model input to a set of NC code and a setup sheet.3.1 Axis of Rotation planningSince the CNC-RP process requires an axis of rotation, the part needs to be analyzed to determine if one exists, and if so, choose the best axis for the process. The axis of rotation can be initially searched based on surface visibility, a necessary condition for machinability. For many parts that would be created with CNC-RP, these axes are most likely the orthogonal axes of the system coordinates. That is, most designers work with a basic 3-view modeling approach,adding features to the views of the orthographic projections of the part (top-front-side sketching in most CAD systems). Similarly, one could argue that most parts intended for manufacturing have been designed for as few setups as possible. So, the first step in the analysis is to simply check the x-, y- and z-axes of the part and determine their visibility.

So, the first step in the analysis is to simply check the x-, y- and z-axes of the part and determine their visibility. In addition, the diameter of the part along each of these axes is found, since smaller diameter materials reduce cost and tooling requirements.(a) Model sliced orthogonal to rotation axisSince tool access is restricted to directions orthogonal to the rotation axis, 2-D visibility maps for a set of cross sections of the surface of the model are used for visibility mapping. This procedure approximates visibility to the entire surface of the model. For example, consider the part illustrated in Figure 3. Cross sectional slices of the geometry from an STL model provide polygonal chains that are used for 2-D visibility mapping. A simultaneous visibility solution for all cross sections of the model will approximate visibility to the entire surface.

A simultaneous visibility solution for all cross sections of the model will approximate visibility to the entire surface. For this simple model and the slice shown in Figure 3a, the chain of edges in the polygon are visible from many different orientations. If the orientations in Figure 3b illustrated by the block arrows are chosen four rotations could be used to machine the part. This implies that four orientations (index rotations) are used and all visible material from each view is removed. If the two orientations noted by the lightning arrows are used, then only two rotations are needed. In this case, two rotations is the fewest number required.(b) Tool access directions restricted to slice planeFigure 3 – Sample model with cross section for visibility mappingto order In this in implement the part CAD/CAM, model is exported as an STL file using a C- program initiated by a The toolbar. custom STL model is then sliced along each of the 3 axes.

custom STL model is then sliced along each of the 3 axes. The visibility analysis is completed for each of the three axes, along with a calculation of the the part diameter of along each axis. This information is displayed to the user as shown in Figure 4. One will note that for this example part, the x-axis is 100% visible and has the smallest stock diameter requirement. The second button in the sequence will then either rotate the model to the chosen axis, or, if none is chosen another algorithm is initiated which will choose the axiszFigure 4 – Initial Visibility analysis of axes for setup orientationyxfor the user. This is a much more advanced option, which uses 3D visibility mapping from the STL file and calculates all feasible axes of rotation. The axis program chooses the best axis, once again based on the smallest stock diameter.

The axis program chooses the best axis, once again based on the smallest stock diameter. The visibility program has been recently published, [L1] and the axis of rotation publication is forthcoming.3.2 Establishing a coordinate systemThe coordinate system for this process must be safely chosen to avoid any chance for tool collisions during processing. Thus, a simple translation of the part is made before processing, to shift the part to the negative x-space and centered on the axis of rotation. This step both ensures that the work coordinate system is consistent for each part and that the part is placed in the center of the stock material. The STL file exported in the previous step reveals the bounding box of the part, which is used to shift the model in the CAD/CAM space. In addition, this approach can ensure collision free processing by establishing a safe working space between the chucks of the fixture system (Figure 5).

In addition, this approach can ensure collision free processing by establishing a safe working space between the chucks of the fixture system (Figure 5). As shown in the figure, the model can be placed in a location such that toolpaths are contained in a safe region (tool collision cannot occur). Moreover, the user will not have to set any tool or work offsets before running a new part.

Moreover, the user will not have to set any tool or work offsets before running a new part. In fact, since the length of the stock is defined in a setup sheet, the tailstock position is set by the length of the stock, therefore no additional checks are required during setup.tailstockindexerchucksDhstockStock length: c = Lp + 2a + 2b where:Lp = Part lengtha = Clamping lengthb = Collision offset (x) = .5Dh +.5 DtMAXbDh = Diameter of tool holder Dt MAX = Diameter of largest toolacFigure 5 – Fixture setup and parameters for coordinate system setting3.3 Adding Sacrificial SupportsA significant challenge in using CNC machining to rapidly create custom shapes lies in the fixturing [W1]. The concept of flexible fixturing has been a popular topic of previous research efforts; unfortunately a completely autonomous fixture design system has yet to be developed [B1].

The concept of flexible fixturing has been a popular topic of previous research efforts; unfortunately a completely autonomous fixture design system has yet to be developed [B1]. The difficult challenge in machining unique part geometries is that the fixtures that hold the material must be generated automatically. The proposed approach to fixturing for this research borrows from the general idea of sacrificial supports, which are used in existing RP systems (i.e. SLA). In this work the general design intent is retained; however, the physical requirements for the support structures are very different. The goal is to have a fixture support solution that is created in-process and is customized for each unique geometry. Specific to rapid machining, the fixture supports need to allow the part to be rotated about the axis while providing access to as much of the surface as possible (Not necessary in additive RP processes such as SLA).

Conventional fixturing methods for CNC often utilize vices, clamps, vacuum surfaces, etc. These approaches occlude visibility to a significant amount of the part or make it difficult to reorient if multiple setups are required. Visibility is so important, simply because it is the necessary condition for being able to cut the surface with a machine tool and to finish cutting the part in as few setup operations as possible.In the proposed method, the sacrificial supports are added to the ends of the CAD model automatically. In this manner, the part remains attached to the stock material throughout the set of machining operations. This leads to better geometric accuracy when we leave the object secured to the parent stock material rather than unclamping, removing, replacing and re- clamping it as in traditional machining setups.

The details of the sacrificial support design methodology cannot be presented in this paper, but the approach can be summarized.In this approach, we consider a worst-case condition for the support design and drive the design based on a maximum allowable deflection. This deflection is considered most significant in the form of twist angle (we approximate with a fixed-fixed statically indeterminate beam model). The design involves a layout of as few as two and as many as four supports in the layout. The design goals are to provide a layout with supports attached 1) near the ends of the part to avoid surface occlusion and 2) far from each other to provide a stiff structure. In order to ensure deflection is limited, we design the diameter for the two “permanent” supports, assuming there are in fact only two supports, and then add up to two more “temporary” supports if possible.

As such, the first two supports are designed based on a simple concentric beam model using a maximum deflection input:4 LrDRD + 1 4 4 4 4 GRrD rLr ) π 1 2 Where: maxθ is derived from a linear tolerance (Rsin maxθ =0.5 tol) R is the radius of the stock material Dt is the largest tool diameter (which determines the lengths of the support) r1 and r2 are the two support diameters2 t4( 4 RrD 2T 2) 4 1tθ=max(++ttThe basic concept is to limit the maximum deflection to at most, half of the required part tolerance. As such, it is assumed that dimensional deviations caused by the support would be the result of under-machining due to deflection of the part The under cutting load. temporary supports simply provide a factor of safety once the permanent supports are this necessary criteria. This of course is a simplification of the entire design model, however, the basic concept is presented.

This of course is a simplification of the entire design model, however, the basic concept is presented. The result of the the design module number, and attachment location of up to four support cylinders. The dimensions are given such that a Chook program in MasterCAM automatically generates the support cylinders. As seen in the figure, the permanent support and temporary supports are different colors. The permanent support will remain throughout the machining process, while the temporary supports can be machined away in a final post processing step. Tests so far have shown that the part dimensions are held with the current support design. Figure 6 illustrates the results of a support design for a ductile iron part requiring 0.006” tolerance. Note that the support lengths in the CAM environment are made arbitrarily long.

Figure 6 illustrates the results of a support design for a ductile iron part requiring 0.006” tolerance. Note that the support lengths in the CAM environment are made arbitrarily long. The actual (machined) length of the supports are determined by the toolpath containment boundaries, which are set based on the largest tool diameter (used in the support equation).designedforisdiameterFigure 6 - Sacrificial support design based on material and tolerance3.4 Orientation and Toolpath planningOnce the sacrificial supports are added to the CAD model, they become part of the drive surfaces for the machining process. The next step is to determine the set of orientations for Rough and Finish machining of the part. In addition, the depths of cut, stepdown, feeds and speeds must be determined for each orientation. is analyzed for visibility about the axis of rotation.

In addition, the depths of cut, stepdown, feeds and speeds must be determined for each orientation. is analyzed for visibility about the axis of rotation. The slice file generated along the axis of rotation yields a set of data that can be used to not only determine part geometry and depths of cut, but to establish the setup orientations. Figure 7 generally represents the feature- free nature of this analysis. As shown, the cross-sectional geometry of the slice file is analyzed to determine the bounds of visibility to the segment’s “left” and “right” indicated by LV and RV. An extensive description of this method is detailed in [F1]. In this step, a Greedy algorithm isRVLVLVRVS1u-1v+S2vu∠vuOnce again,the modelFigure 7 – Feature-free visibility and setup calculations using cross sectional geometryuv∠used to quickly find a set cover for the angles of orientation required to machine the part.

The maximum visible orientation angle is the seed value for an initial roughing operation, which consists of mass material removal to remove excess stock material about the part. Next, a rough machining of the part surfaces using a large stepdown (~0.030”) is created. The visibility algorithm determines the first and last visible surfaces on the part, from each orientation. The first mass material removal operations machines from the stock radius to the first visible surface of the part from that setup angle, then the rough machining operation cuts from the first part surface to the furthest visible surface. Figure 8 below illustrates the depths used in rough machining. The roughing toolpaths are generated using a heuristic method developed from numerous part trials. The goal of the roughing operations is simple; remove as much stock material as possible, while avoiding thin webs of stock material that may cause catastrophic failure.

The goal of the roughing operations is simple; remove as much stock material as possible, while avoiding thin webs of stock material that may cause catastrophic failure. The problem is that during a set of rotated toolpaths, one may find that the tool is being forced to cut through the back facing surface, which may be a thin web of material. During this process, the thin material deflects and sometimes can wrap around the tool and cause failure. The roughing routines are therefore sequenced to a specific order of setup orientations such that the stock material is removed in a controlled, incremental manner.

The roughing routines are therefore sequenced to a specific order of setup orientations such that the stock material is removed in a controlled, incremental manner. Figure 9 illustrates the methodology behind thin web avoidance, used to create 3 initial roughing paths to machine most of the stock safely from the part.Mass removalSecond roughing toolpathsFigure 8 – First two roughing operations for removing stock from round materialTool selection in a traditional machining process typically involves feature-specific selection whereby the tool geometry is based on the features to be created (such as a drill for a hole). rapid prototyping environment, we typically use one or very few “tool” settings, such as the laser spot diameter in SLA, SLS or tip diameter in FDM. As such, the tool is intended to be generic, and generates shapes iteratively, such as hatching patterns with laser. Similarly, the CNC-RP method uses a general and robust strategy.

As such, the tool is intended to be generic, and generates shapes iteratively, such as hatching patterns with laser. Similarly, the CNC-RP method uses a general and robust strategy. Tools are chosen base on their necessary criteria; long enough to reach the part surfaces. Thus, a simple tool library contains sets of tools with increasing diameter. For diameter, the tool in the library is as long as possible/available. Therefore, when aPart geometryExcess stock(a)(b)(c)However,inaFigure 9 – Roughing strategy (a) part is initially machined from highest visible angle, then (b) and (c) orientations remove stock while avoiding thin web conditionsthedepth of cut is determined based on the visibility algorithm, the tool chosen will have a minimum diameter. Although this may not be efficient, it is necessary for a truly feature-free process planner.

Although this may not be efficient, it is necessary for a truly feature-free process planner. That is, if we assume nothing about the surface geometry, then we must assume that small features might be present on the surface, hence, a small diameter tool must be chosen. Then, tool parameters are very straightforward, such as feeds, speeds and stepdowns. The feeds and speeds are essentially the maximum that are allowable for each material type, given the stepdown amount. The stepdown in CNC-RP is actually analogous to the “layer depth” in a traditional additive RP process. So, the tool removes material layer by layer (for each stepdown), rather than adding material (in most CNC-RP toolpaths, this layer depth is 0.020” to 0.050” for roughing and 0.001”to 0.003” for finishing).Finally a set of finishing toolpaths is developed.

As for the second roughing operations, the finishing toolpaths begin the first visible surface, cutting at however, the depth of the final toolpath layer is determined using a simple optimization routine. For the set of setup orientation angles defined for the part, the visibility algorithm sorts the surfaces to be machined by allocating each surface to the setup angle that can surface with minimum reach distance. As such, the set of finishing toolpaths do not need to machine to the furthest visible surface from that angle, rather, they must machine to the furthest visible surface for all surfaces that are closest to stock radius at that angle. This is calculated using depth calculations from the slice geometry of the part file. Figure 10 illustrates the optimization using a circle representing the sectional “slice”. It illustrates three setup orientation angles about the axis the of rotation and comparison a of deepest-visible versus minimum-distance allocation method.

It illustrates three setup orientation angles about the axis the of rotation and comparison a of deepest-visible versus minimum-distance allocation method. As(a)(b)Figure 10 – Toolpath depth planning optimization; (a) deepest visible, (b) minimum distancethecrossan implementation in the CAD/CAM system, some user defined are solutions permissible. Although the CNC-RP system final the makes determination on roughing angles andFigure 11 – User interface for setup orientations and system defined solutionssettings, the user can force certain setup directions into the finishing set. Therefore, the greedy algorithm will consider the user’s input as forced solutions, and then optimize based on that reduced set (those surfaces in the set not covered by the user-defined angles). This method is employed because the user may wish to select some angle (perhaps one that aligns with a desired feature) that they believe is a required or good angle.

This method is employed because the user may wish to select some angle (perhaps one that aligns with a desired feature) that they believe is a required or good angle. Figure 11 illustrates a simple user interface, and then the solution as derived by the system. In this case, the user forces 0 and 180 degrees into the solution, then the greedy algorithm finds three additional angles for the set cover.It should be noted that the toolpaths generated for both rough and finish machining are generated from the original CAD model of the part, as imported into the CAD/CAM system. Therefore, the toolpaths in CNC-RP are not based on an STL file, rather, from the native surface geometry. Although STL models of the part are used throughout the analysis of the geometry (for setup angles, supports and toolpath parameters), the actually toolpath generation programs are always run on the accurate part surface.

In this manner, CNC-RP avoids the inherent approximation errors of additive processes which derive their toolpaths from the STL model.4. Examplesthe process steps of described in the previous section result in a concatenated set of NC code and a setup sheet. The setup sheet lists the tools required and their tool changer locations, and the diameter and length of the stock material. In order to run the part program, the user simply loads the material and NC code an initiates a cycle-start.TheFigure 12 – Femur bone fixtured via sacrificial supportsFigure 12 illustrates an entire femur bone near completion in a rapid machining process,fixtured using sacrificial supports. This bone was finish machined from 3 orientations, as determined by the geometric algorithms in approximately 8 hours in Delrin plastic.step-by-step illustrates machining of a cast iron component from a 5” diameter bar stock.

The machining time for this part was approximately 24 hours, however, the NC code was generated from a CAD model in under 5 minutes. In both the laboratory and an industrial sponsor installation, the CNC-RP process has been shown to effectively create a variety of parts using functional materials from Plastic, ranging Aluminum, Cast Iron and Steels.Figure 13theFigure 13 – Example steps during the rapid machining of a cast iron component5. ConclusionThe implementation in a CAD/CAM environment enables CNC-RP as an effective rapid prototyping method for use on existing CNC machines. As such, the process can be used for short run production, custom manufacturing and prototyping for certain applications. In addition, the same CNC machine can also be free to create production parts in conventional applications. The impact is that there is no need for a specialized RP machine for these functional prototypes and parts.

The impact is that there is no need for a specialized RP machine for these functional prototypes and parts. The system is currently being tested as a method for creating spare parts for legacy equipment in the agricultural industry. Of course, the process has limitations in both geometry and scale. Traditional additive RP processes are infinitely capable at creating complex shapes, as compared to CNC machining. In contrast, however, CNC-RP is generally more capable in surface finish, since layer depths can be controlled to very small values, as compared to additive RP processes. The greatest advantage however, is that CNC RP has exceedingly better capability in using a variety of materials to create truly functional parts.B1. Bi, Z.M. and Zhang, W.J., “Flexible fixture design and automation: Review, issues and future directions”, International Journal of Production Research, Vol. 39, No.13, pp. 2867-2894, September 2001C1.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIMPLICIT SLICING METHOD FOR ADDITIVE MANUFACTURING PROCESSESD. W. Adams, C. J. Turner1Department of Mechanical Engineering, Clemson University, Clemson, SC 29631AbstractAll additive manufacturing processes involve a distinct preprocessing stage in which a set of instructions, or GCode, that control the process specific manufacturing tool are generated, otherwise known as slicing. In regards to fused deposition modeling, the GCode defines many crucial parameters which are needed to produce a part, including tool path, infill density, layer height, feed rate, tool head and plate temperature.

The majority of current commercial slicing programs generate tool paths explicitly, and do not consider particular geometric properties of parts, such as thin walls, small corners and round profiles that can result in critical voids leading to part failure. This work replicates an implicit slicing algorithm in which functionally derived infill patterns are overlaid onto each layer of a part, reducing the possibility of undesired voids and flaws. This work also further investigates the effects that varying implicitly derived infill patterns have on a part’s mechanical properties through tensile testing dog bone specimens with three different infill patterns and comparing ultimate stress and elastic modulus properties.1.

IntroductionAlthough additive manufacturing (AM) has been in practice since the mid 1980’s, it has recently become a more common method of mass part production due to time and cost savings, freedom of design, and the ease of implementation it provides in comparison to other methods of manufacturing [1]. The transition that additive manufacturing has made from being mostly a small market DIY part prototyping fad, to a legitimate industry, capable of providing solutions to previous manufacturing challenges, can be largely attributed to the numerous studies which have been completed to advance our knowledge and perception on this topic.The additive manufacturing industry encompasses several different methods of rapid prototyping including Sterolithography [2], Electron Beam Melting (EBM) [3], Direct Metal Deposition (DMD) [4-5], Selective Laser Sintering [6-8], and Fused Deposition Modeling (FDM).

While all of these processes offer unique advantages for particular prototyping applications, they all implement the same preprocessing method known as slicing. A slicer algorithm takes a three- dimensional model and generates a set of basic instructions, known as GCode. This GCode controls every build parameter and every move that the respective manufacturing tool makes throughout the entire build [9]. Slicers work by intersecting the input model with distanced parallel planes to create a set of two dimensional domains; that, when stacked on top of each other, represent the entire model as seen in figure 1. A tool path is then created for each “slice”, and the part is manufactured layer by layer, from the bottom up.1 Corresponding Author, E: cturne9@clemson.edu844Figure 1. Perimeter curves and Delaunay Triangulation of all slices of dog bone specimenCurrent commercial slicer algorithms generate the tool path for each slice explicitly.

Perimeter curves and Delaunay Triangulation of all slices of dog bone specimenCurrent commercial slicer algorithms generate the tool path for each slice explicitly. In other words, the tool path is generated by applying a series of predetermined geometric transformations to the interior of each slice without considering overall part geometry. Tool path predetermination can be problematic in various geometric situations, such as thin walls, small corners and round profiles, and parts can be left with critical voids that may lead to failure. Work has been done in developing alternate slicing methods which take part geometry into account to not only solve this issue but also to optimize different features of the slicing and printing process. This work replicates and further investigates an alternative implicit slicing method for fused deposition modeling in an attempt to solve the unfilled region issue that the explicit method presents [10-12].

The implicit algorithm overlays a user defined functionally derived infill pattern onto the intersection perimeter for each slice ensuring that no voids will occur. Additionally, not only does this allow for a much wider array of possible infill patterns, it also allows the manufacturer to apply particular infill patterns that are better suited for an individual part based on its application. This work also experimentally examines the effects that varying implicit infill patterns have on the mechanical properties of a part.While this work does conclude that an implicit slicing method offers a concrete solution to part voids, the current state of the current work does not confirm that implicitly derived infill patterns offer more ideal mechanical properties than explicitly derived infill patterns. Although Stueben et al. state there are material property benefits with implicit infill patterns, further testing to compare the two methods needs to be performed to confirm this claim.2.

Although Stueben et al. state there are material property benefits with implicit infill patterns, further testing to compare the two methods needs to be performed to confirm this claim.2. Background and Related Work2.1 Fused Deposition ModelingFused Deposition Modeling (FDM) is an AM process in which an extruder tool head follows a coded path while extruding a specific amount of thermoplastic filament layer by layer to form a complete part. The thermal properties of the thermoplastic in use allow each layer of the part to bond together which allow the resulting part to be considered one continuous entity. Figure 2 displays the basic process of fused deposition modeling.845Figure 2. FDM Process [13]A combination of several stepper motors control the x and y locations of the tool head during the build, which are contingent on the infill pattern that is created for each layer.

FDM Process [13]A combination of several stepper motors control the x and y locations of the tool head during the build, which are contingent on the infill pattern that is created for each layer. After one layer is completed, a lead screw lowers the build plate a user defined distance so the next layer can be extruded on top of the previous layer. The thermoplastic is heated, and remains in a molten state long enough to allow complete bonding between the layers after extrusion. In particular cases, as the one pictured in figure 1, parts may require support material to ensure that none of the extruded material collapses, causing a build failure. Once the build is complete, the support material is removed by either breaking it off or dissolving it in a material specific solvent.There are many different user defined build process parameters that have been found to significantly affect the overall quality of a part.

Some of these parameters include layer height, extrusion resolution, spindle speed, feed rate, part orientation, retraction distance and rate, support material settings, infill pattern, and infill density. All of these parameters collectively affect build properties such as surface roughness, build time, amount of extruded material, and the mechanical properties of the part; therefore, they should all be carefully considered before printing to ensure optimal manufacturability and a quality finished part.Many studies have been completed to exemplify the impact that varying explicitly derived infill patterns have on material properties, such as ultimate strength [14], strain rate, modulus of elasticity and yield strength. One important study that was done by Ziemian et al. experimentally validated the idea that raster orientation significantly affects all of the mechanical properties just listed above [15].

One important study that was done by Ziemian et al. experimentally validated the idea that raster orientation significantly affects all of the mechanical properties just listed above [15]. The study implemented four different raster angles, longitudinal (0°), diagonal (45°), transverse (90°), and an alternating +45°/-45°. Results concluded that when the extruded filament was aligned with the direction of the force being applied, the part exhibited much higher values for all mechanical properties including a 35% higher yield strength, 33% higher ultimate strength, and 29% higher elastic modulus than any other test angle.2.2 SlicingAs previously mentioned, every additive manufacturing process undergoes the same preprocessing procedure known as slicing. All slicing programs start with a 3D CAD (Computer Aided Design) model and transform it into a script that the AM machine’s processor can understand and implement to create a working part.

All slicing programs start with a 3D CAD (Computer Aided Design) model and transform it into a script that the AM machine’s processor can understand and implement to create a working part. Currently the vast majority of commercial and industrial slicer algorithms operate using the same accepted approach, which is considered to be an explicit slicing method. The main defining characteristic of an explicit slicer is that the interior tool path, or infill pattern, for each layer is generated the same way for every single input model, with no regard to the perimeter geometry. The infill path begins at an arbitrary point along the perimeter of the slice, usually close to the ending point of the perimeter, and starts to apply a846discrete set of geometric transformations to that path in order to create the desired pattern. This process is done for every layer, excluding the exterior layers of the part, until the part is complete.

This process is done for every layer, excluding the exterior layers of the part, until the part is complete. This slicing method has proven to present an issue with parts containing critical voids, due to an incompatibility of the input geometry with the geometric transformations, such as thin walls and small corners. Figures 3 and 4 displays an example of this problem.Figure 4. Discontinuity between round profile and infillFigure 4. Model exemplifying lack of infill in thin wallsSlicers that explicitly compute tool paths usually offer 5-10 preset infill patterns, with common ones being a honeycomb, linear, rectilinear, diamond, triangle, and concentric.

Due to the limited amount of options for infill patterns, users are unable to tailor the infill to a part’s specific needs based on its application.Multiple experimental slicer algorithms have been developed to address both of theseissues that explicit slicers present as well as optimizing different aspects of the slicing process.One alternative slicing algorithm, developed by Lu et al., considered object breakability as the backbone of the methodology [16]. The team of researchers drew inspiration from the Voronoi structure to allow printed material to be distributed unevenly. After a FEM stress simulation was implemented on a given model and the data was extracted, the algorithm runs through a series of optimization processes that calculate weight to strength ratios for several nodes that were derived from a density map defined by the stress distribution.

This process places more Voronoi cells in areas where the stress is higher so that the extruded material is concentrated in locations where the model is most likely to break. The tool path is then created based on the contours of all of the cells. This method proved to reduce material volume by up to 75% while still maintaining the structural integrity of the model.Micali and Dornfeld developed a three-dimensional slicer for point-based AM systems that enable tool paths to be generated in three dimensions, rather than a planar surface, to preserve as much geometric continuity of the model as possible [17]. It considers the opportunity of improvement for AM machines that have 5+ degrees of freedom (DOF) instead of the standard 2.5 DOF for commercial machines.

It considers the opportunity of improvement for AM machines that have 5+ degrees of freedom (DOF) instead of the standard 2.5 DOF for commercial machines. This method would allow much higher precision and accuracy for machines in the attempt to make a prototyped part as similar to the design as possible.In an effort to significantly reduce prefabrication computation time for the slicing method and mass customization, Ye et al. developed an alternative slicing algorithm that took advantage of the ability to reuse topology information [18].

developed an alternative slicing algorithm that took advantage of the ability to reuse topology information [18]. It was created in order to deliver parts that could more precisely meet individual needs for a similar, yet changing, model without losing efficiency.847The team proposed to store topological information at three different levels of the prefabrication process, allowing tool paths to be generated for slightly changed models in significantly less time, rather than completely recalculating a tool path every time the model is slightly changed.Another significant issue with conventional explicit slicing is the lack of manufacturability consideration. Tedia and Williams used a voxel-based geometric modeling method to optimize part orientation prior to being prototyped [19]. This was done by defining specific exterior voxels as support voxels based on the angle that their normal makes with the horizontal.

This was done by defining specific exterior voxels as support voxels based on the angle that their normal makes with the horizontal. An optimization process is then implemented to maintain that the minimum number of support voxels exist. Their findings proved to minimize necessary support material to maintain structural integrity as well as improve time of manufacturing.The purpose of this work is to replicate and optimize an implicit slicing method that was proposed by Stueben et al. This work attempts to address the issue of critical voids in additively manufactured parts, in which explicit methods present, as well as the issue of infill and mass customization. Several infill patterns are generated with this method, and three test cases are prototyped and tensile tested to investigate the effect that varying implicit infill patterns have on mechanical properties.3.

Methodology3.1 Slicer DevelopmentThe first process that must occur before a model can be sliced is the transformation of a 3D model into a STL file. A STL file is a set of points in the three-dimensional Cartesian coordinate system that represent nodes of all of the triangles which the model is spliced into when it is meshed, and ultimately the entire model when put together. This process is required so that a computational language, in this case Matlab, can interpret and display the model. After the coordinates are read into a script, the model can then be sliced, and a tool path can be generated. The first technical step that every planar slicing algorithm implements is finding intersection curves of the model, so that a perimeter curve can be defined for each slice. To do this, parallel planes, which are intersected with the part model and point where the planes intersect the part, are stored and a perimeter curve is created for that one slice.

To do this, parallel planes, which are intersected with the part model and point where the planes intersect the part, are stored and a perimeter curve is created for that one slice. This process is iterated for a calculated number of times, based on the overall part height divided by desired layer thickness. If meshed together, this discrete set of perimeter curves will represent the entire model. For this experiment, a dog bone specimen of 3.0mm thickness was chosen as the base model based on its common use in tensile testing. A layer thickness of 0.2mm was chosen; therefore, 15 layers would be created for this model. Figures 5 and 6 depict this perimeter declaration process with the dog bone specimen.848Figure 6. Plane intersection of modelFigure 6. Intersection points resulting from intersection of plane and modelAfter the outer perimeter curves have been declared, a certain amount of interior offset curves need to be created, based on user definition.

Interior curves are necessary to ensure bonding from the infill pattern to the boundary curve, as well as to create a stronger shell for the part. There are many diverse methods for offsetting polygons such as the use of Voronoi diagrams [20, 21]; however, this work implements a signed Euclidean distance function to obtain equidistant interior offset curves [22]. Results of this transform are displayed in figures 7 and 8.Figure 8. Contour plot of Euclidean distance functionFigure 8. Interior shell contour with an offset distance of .3mmThe last step before being able to apply an infill contour to a layer of the model is to define a domain for each slice. The domain represents all of the available space that an infill curve can occupy, ensuring that the infill pattern does not overlap or exit the perimeter contours.At this point, any mathematical function can be applied to the domain of a layer to create infill contours.

In essence, a function is being overlaid onto the domain of the slice; therefore, no region of the model will be void due to geometrical complications. After functions are applied to the domain, contour plots with a specific number of contours can be generated, so that the amount of extruded material and overall part mass can be controlled. The points that define each contour are concatenated with each other and then with the set of perimeter curve points in order to make a consolidated list of all the points that the tool head will extrude material over, thus completing the infill pattern for one layer. This process is done for every layer of the model, thus creating a list of every point along the tool path for the entire build. Figure 9 displays various possible functionally derived infill patterns.The final step of the prefabrication phase of a slicing program is GCode generation.

Figure 9 displays various possible functionally derived infill patterns.The final step of the prefabrication phase of a slicing program is GCode generation. All AM machines require a set of instructions that will control every setting and move that the respective tool head will make. This set of instructions is known as GCode. Not only does the GCode script contain the coordinates for the tool path, but it also defines certain print settings that are critical to a successful prototype, including temperature of the tool head and platform, raft settings, tool head speed, retraction distance, and various other preprint procedures that are defined849by the user. After the code is formatted and compiled, it is sent to the particular AM machine and the part can then be manufactured.Figure 9.

After the code is formatted and compiled, it is sent to the particular AM machine and the part can then be manufactured.Figure 9. Possible examples of functionally derived infills that can be applied using implicit slicing method.To test the validity of the application of this implicit slicing method, as well as investigate the effects of varying implicitly derived infill patterns on mechanical properties, three test patterns were chosen to be prototyped and tensile tested. Mechanical properties including ultimate strength and modulus of elasticity were analyzed.3.2 Experimental SetupAll of the dog bone specimens manufactured for this study were made from acrylonitrile butadiene styrene (ABS) by a Makerbot Replicator 2X 3D printer. The Replicator 2X is equipped with two extrusion heads, a heated glass platform, and a wide array of customizable parameters so that each print can be altered for a part’s specific needs. ABS was chosen as the material for thisFigure 10.

ABS was chosen as the material for thisFigure 10. Experimental testing equipment and set up with psylotech µTS 10kN tensile tester850work based on its optimal thermal properties, allowing a secure cohesion to the printing surface and maximum bonding between layers.After being manufactured, every test specimen was tensile tested using a psylotech µTS 10kN tensile tester (pictured in figure 10) with friction grips placed 18mm from the edge of the part. The implemented testing profile started at 0N and pulled at a constant distance rate of 75µm/s. The testing software recorded displacement and force data as the part was being pulled at a rate of 100Hz.

The testing software recorded displacement and force data as the part was being pulled at a rate of 100Hz. Stress and strain data were then extrapolated from the recorded data using mathematical relationships.3.3 Test CasesThe first implicitly derived infill pattern chosen as a test case was the diagonal hatchpattern, which was computed with[1](𝑧𝑧 𝑙𝑙𝑡𝑡⁄ )%2𝐻𝐻𝑙𝑙𝑙𝑙𝑙𝑙(𝑥𝑥, 𝑦𝑦, 𝑧𝑧, 𝜃𝜃) = 𝑥𝑥 𝑠𝑠𝑠𝑠𝑠𝑠(𝜃𝜃) + 𝑦𝑦 𝑐𝑐𝑐𝑐𝑠𝑠(𝜃𝜃)(−1 where x, y, and z are the Cartesian coordinates of the dog bone, l t is the layer thickness, θ is the amount each line is rotated from the x axis, k is a constant that allows the amount of extruded material (mass) to be controlled, and % represents the modulo operator. This diagonal hatch pattern was chosen due to its widespread use in commercial additive manufacturing. This pattern is displayed in figure 11.) 𝑘𝑘The second infill pattern chosen was the egg crate pattern. The egg crate pattern was chosenFigure 11.

This pattern is displayed in figure 11.) 𝑘𝑘The second infill pattern chosen was the egg crate pattern. The egg crate pattern was chosenFigure 11. Hlin(x,y,z,π/4) – Diagonal Hatchto mimic a mechanical behavior similar to that of a honeycomb style pattern. Honeycomb patterns are also widely used in industry due to its favorable strength to volume ratio; however, explicit slicers often have trouble effectively filling thin walled structures with this pattern, as seen in figure 4. The explicit egg crate pattern was calculated with[2]22 + sin(𝑏𝑏𝑦𝑦)𝐻𝐻𝑒𝑒𝑒𝑒𝑒𝑒(𝑥𝑥, 𝑦𝑦, 𝑎𝑎, 𝑏𝑏) = sin(𝑎𝑎𝑥𝑥) where x and y are the planar Cartesian coordinates of the dog bone, and a and b are arbitrary constants that can be altered to change the density of the pattern. This pattern is displayed in figure 12.851Figure 12.

This pattern is displayed in figure 12.851Figure 12. Hegg(x,y,.43,.43) - Egg crate infill pattern.The final test pattern was generated in part by using contours from a stress distribution, which was obtained by running a FEA simulation on the dog bone specimen using SolidWorks. The force that was applied during the simulation was the same as it would be during actual tensile testing. Based on the fact that material fibers exhibit maximum resistance to failure when force is applied along the same axis as the fiber, (i.e. if the filament was aligned with the stress contours and was more concentrated in regions that experience more stress), the specimen would prove to be stronger and have different failure tendencies. To simulate the same forces that the dog bone would undergo during testing, a force of 125N was applied to the top and bottom faces of the left side of the dog bone with a grip length of 18mm while the right side of the model was fixed.

A stress distribution was then obtained from the results of the simulation.Having an infill pattern comprised of only the stress contours would be insufficient due to resulting large gaps in the model; therefore, modifications must be made. The stress derived infill pattern was computed with𝐻𝐻𝑙𝑙𝑙𝑙 = 𝜎𝜎𝑣𝑣𝑣𝑣(𝑥𝑥, 𝑦𝑦, 𝑧𝑧)𝐻𝐻𝑙𝑙𝑙𝑙𝑙𝑙(𝑥𝑥, 𝑦𝑦, 𝑧𝑧, 𝜋𝜋 4⁄ ) 𝑘𝑘 where σvm(x,y,z) is the three dimensional mesh of the stress values post simulation. The combination of the stress distribution and the linear hatch contour plots concentrate more contours in regions that experience the highest stress values, as mentioned previously. Figure 13 displays the contour plots that comprise the final stress derived infill pattern, which can be seen in figure 14.4. Results4.1 Test Case 1 – Diagonal HatchUpon testing, all specimens with a diagonal hitch infill pattern failed in the neck region; however, the exact failure location was not consistent.

Results4.1 Test Case 1 – Diagonal HatchUpon testing, all specimens with a diagonal hitch infill pattern failed in the neck region; however, the exact failure location was not consistent. Failure location was evenly spread out along the neck and no failure trend was apparent. Further visual inspection indicated that the failure was a very brittle, uneven break into two distinct parts of varying proportion. Additionally, specimens with the diagonal hatch pattern displayed deformation behavior similar to that of a solid sample, as most deformation occurred on either side of the failure.852[3]abcdeFigure 13.

(a) Stress derived contour plot (b) Diagonal hatch for odd numbered layers (c) Infill pattern for odd numbered layers (d) Diagonal hatch for even numbered layers (e) Infill pattern for even numbered layersFigure 14.Stress derived infill pattern.𝐻𝐻𝑙𝑙𝑙𝑙 = 𝜎𝜎𝑣𝑣𝑣𝑣(𝑥𝑥, 𝑦𝑦, 𝑧𝑧)𝐻𝐻𝑙𝑙𝑙𝑙𝑙𝑙(𝑥𝑥, 𝑦𝑦, 𝑧𝑧, 𝜋𝜋 4⁄ ) 𝑘𝑘The total amount of extruded material for each print in this test case was 1348.78mm. Samples for this test case had an average mass of 3.197 grams and an average print time of 45 minutes. Regarding mechanical properties for the diagonal hatch infill pattern, the average ultimate stress was 14.6 MPa, and the average elastic modules was 1.08 GPa.4.2 Test Case 2 – Egg CrateSpecimens with the egg crate infill pattern all failed in the transition region that lies between the grip and neck regions. Visually, the failures can be described as a very sudden, brittle fracture.

Visually, the failures can be described as a very sudden, brittle fracture. After inspection, the bulk of deformation occurred at the fracture point, while the rest of the specimen underwent very little deformation.The total amount of extruded material for the egg crate infill pattern was 1384.49mm. Samples for this test case had an average print time of 44 minutes and an average mass of 3.013 grams. The mean ultimate stress was the minimum out of all three test cases at 14.3 MPa. The average elastic modulus was less than that of the stress derived test case but still more than the diagonal hatch infill pattern with a value of 1.14 GPa.4.3 Test Case 3 – Stress Derived PatternFailures for specimens with the stress derived infill pattern were much more inconsistent than those in test cases 1 and 2. All of the failures began with a small crack on the perimeter of the sample, which eventually propagated.

All of the failures began with a small crack on the perimeter of the sample, which eventually propagated. There was an even distribution of samples which failed due to fatal deformation at the point where the grips clamp the sample and failures that were located at the transition region of the dog bone.853The amount of extruded material for the stress derived infill pattern was 1352.13mm. Samples for this test case had an average print time of 43 minutes and an average mass of 3.135 grams. This mass is less than the average mass noted in test case 1, which can be attributed to the complexity of the pattern. The average ultimate stress for this test case was 15.9 MPa, and the average modulus of elasticity was 1.31 GPa.4.3 AnalysisA table comparing the nominal values of all measured variables is displayed in Table 1.Table 1.

Table comparing the experimental values for all 3 test casesThe independent variables part geometry, mass, and extruded filament were all held at a constant within a small margin of error to exemplify the effects that varying infill patterns have on mechanical properties. Quantitatively, the stress derived infill pattern exhibited the largest value for ultimate stress the other test cases by roughly ~10%. In regards to the elastic moduli of all three infill patterns: none of the test cases exhibited strain values greater than 2%. Theoretically, this indicates that, on average, none of the specimens entered an inelastic state. In this case, the average modulus of each test case was calculated by creating a linear line of best fit for each data set. The stress derived infill patterns demonstrated an average modulus ~20% greater than that of the diagonal hatch pattern and ~6% greater than that of the egg crate pattern.

The stress derived infill patterns demonstrated an average modulus ~20% greater than that of the diagonal hatch pattern and ~6% greater than that of the egg crate pattern. This supports the claims hypothesized based on the alignment of the extruded filament with stress contours. Figure 15 displays average stress vs. strain curves with lines of best fit for each test case.854Figure 15. Stress vs. Strain Curve of all three test cases with Elastic Modulus equations labeledAll of the specimens from each different test case also exhibited dissimilar failure locations. Specimens with a diagonal hatch infill pattern all fractured and ultimately failed in the neck region, which is consistent with the failure location of a continuous dog bone sample. Specimens with egg crate and stress derived infill patterns exhibited failures that would be considered abnormal.

Specimens with egg crate and stress derived infill patterns exhibited failures that would be considered abnormal. Every egg crate specimen failed due to fractures in the transition region of the dog bone, and almost half of the stress derived specimens failed due to fractures adjacent to the grips. This phenomenon is important to note because failure location has previously been predicted by geometry and material type alone; however, it is clear that infill pattern has a large impact on point of failure for additively manufactured parts.5. Conclusions and Future WorkThe recent proliferation of additively manufactured parts can be attributed to many different factors, including time and cost savings, as well as improved manufacturability. As more additively manufactured parts are used in applications, in which loads are applied, it will become increasingly important to know how the respective parts will react, as well as, do everything possible to optimize the process.

The purpose of this work was to address some of the issues that the explicit slicing method presents, through the development of an alternative implicit slicing method. This slicing method ensures that no voids would be present in tool paths based on input geometry and offers a much wider array of possible infill patterns than the explicit method.Additionally, based on the results of experimental testing, it is clear that all of the implemented implicitly derived infill patterns affected the part’s mechanical properties differently. The stress derived infill pattern displayed the highest values for both ultimate stress and average modulus of elasticity, with values of ~10% and ~21% greater than those of the diagonal hatch855specimens respectively.

Gathering experimental data of mechanical properties for a statistically significant sample of infill patterns would allow an appropriate implicitly derived infill pattern to be automatically generated, based on input geometry, desired mechanical properties under a particular load, and failure location.While there is currently a significant amount of work being done in attempts to optimize infill patterns within the AM industry, there is still much that can be done. The most advantageous next step, beyond this work, would be to prototype and mechanically test additional infill patterns to see how they behave when subjected to different loading, including bending, torsion, and compression. If a statistically significant amount of data was collected for many different infill patterns under different loads, neural networks could be formed and the optimal infill pattern could be determined.

Ideally, instead of the user arbitrarily defining an infill pattern, it would be automatically selected based on desired mechanical properties and input geometry. Additionally, applying this slicer algorithm and testing procedure to additively manufactured metal parts would be beneficial. Another study which could further this work would be the development of a mathematical function that could be implemented with the implicit slicer which would control density of the part. Being able to control the density throughout a part would ultimately allow you to control the failure location, which could be useful in certain applications, such as shear pins.ACKNOWLEDGEMENTSThe authors would like to acknowledge the support of Clemson University and the Naval Research Laboratory in this work.

PROVED ENERGY DELIVERY FOR SELECTIVE LASER SINTERINGCarl Deckard and David MillerClemson University Clemson, SC 29634ABSTRACTSelective Laser Sintering (SLS) is a leadingtechnology in the important new area of Solid Freeform Fabrication (also ca1ledR.apid Prototyping). Selective Laser Sintering produces freeform parts directly from a CAD model by building the parts up in layers from a powder. A laser is used to selectively melt each layer of powder to form the part. The laser beam is scanned across the powder using two galvanometer scanners. The energy delivery system (laser, optics, scanner and controls) is a critical component technology of SLS. Projects with the objective of improving the energy delivery system are underway at Clemson University.INTRODUCTIONSelective Laser Sintering (SLS) is a leading process in the new field of Solid Freeform Fabrication (SFF).

SLS produces parts directly from a CAD model by melting or sintering thin layers of a powder together with a laser beam. SLS is a thermal process that is applicable to a wide range of materials.At Clemson, we are working on SLS with the objective of developing process understanding that will lead to system performance improvements. We are concentrating most of the research effort on the energy delivery system (laser, optics and scanner). Our near term focusing objective is to improve the part production speed. Control of the thermal gradient within the layer being sintered (or melted) has been identified as a key process issue.BACKGROUNDSLS and other SFF processes have the potential to become indispensable, mainstream industrial processes. At this time, SLS and other SFF processes are primarily used by large companies and service bureaus to make small production number high value parts. To extend SLS and other SFF processes to new applications and users, continuous ........

To extend SLS and other SFF processes to new applications and users, continuous ........ 11"'......... ',,"" ........ "",1'\1" in process performance is needed. Several key performance measures are: accuracy, 2) material properties, 3) part production speed, 4) surface finish, 5) part production cost, and 6) maximum part size. Improvements in several areas are needed to achieve these performance improvements. These areas include: materials, machines (hardware and controls) and software.Statement of Need2.2 Current Energy Delivery System The energy delivery system hardware currently used in commercial SLS machines consists of a laser, focusing optics, and a scanning system.151The laser is a commercially available 50 watt C02 laser. A two lens system is used to focus the beam. The scanning system consists of two mirrors which are rotated by galvanometers to deflect the laser beam to the desired position. The galvanometers are controlled by an analog PID loop and a servo amp.

The galvanometers are controlled by an analog PID loop and a servo amp. The set points are generated by the scan computer in response to vector data provided by the process control computer.The scan pattern generally consists of an outline and a parallel line fill pattern (Fig. 1). The spacing between the scan lines is generally one half to one quarter of the beam diameter; therefore, each point to be scanned is hit by the laser beam two or more times (Fig. 2). The laser beam spot size is generally around 0.4 mm. The minimum attainable laser beam spot size is governed by diffraction of the laser beam and the maximum desirable spot size is set by the need to produce parts with fine features.2.3 Laser Material Interaction and Heat TransferSome of the beam energy is reflected by the powder and the rest is absorbed by the powder. The depth of penetration of the laser beam into the powder bed is a function of the optical properties of the powder at the laser wave length (1).

The depth of penetration of the laser beam into the powder bed is a function of the optical properties of the powder at the laser wave length (1). See Figure 3. The intensity of the laser beam within the powder bed as a function of depth is:I(z) = (1- R)Ioexp(-bz)Where:R = surface reflectivity10 = laser light intensity at the surface~ = extinction coefficientz = depth into the bedThe absorption of the beam by the powder creates a three dimensional heating of the powder. The heat input distribution is a decaying exponential multiplied by the spot profile (usually a gaussian). The energy input at a point is time varying due to the motion of the beam.

The heat input distribution is a decaying exponential multiplied by the spot profile (usually a gaussian). The energy input at a point is time varying due to the motion of the beam. From reference (1), the heat source function is:Where:x,y= horizontal position of the point in questiong(x,y,z, t)= heat input as a function of position and timet= timew= spot sizev x= beam velocity in the x directionv y= beam velocity in the y direction152(1)(2)The resulting temperature distribution is the initial condition for a three dimensional transient heat transfer problem with variable coefficients (1). The heat transfer is described by the conduction equation:aT'pCat =V(kVT)+ g(x,y,z,t)Where:p = density C = specific heat k = thermal conductivityThe density and the thermal conductivity change with degree of sintering. Radiation and convection from the surface must also be considered. This problem has been dealt with in great detail by other authors (1).

Radiation and convection from the surface must also be considered. This problem has been dealt with in great detail by other authors (1). The full results are not presented here but some general conclusions should be noted. If the laser beam interaction time is small with respect to the time constant for heat transfer, then the thermal gradient is primarily a function of the average temperature rise of the powder layer due to the one pass of the laser and the extinction coefficient. If the laser beam interaction time is on the same order or longer than the time constant for heat transfer, then the thermal gradient will be lower than predicted above due to heat transfer during the laser material interaction.2.4 Significance of Thermal GradientsLaser material processing is normally characterized by larger heat fluxes and as a result, large thermal gradients. Large thermal gradients are a significant feature of the SLS process.

Large thermal gradients are a significant feature of the SLS process. The acceptable numerical range for the thermal gradient is bounded. If the thermal gradient is too low, then the part will not have well defined edges. If the thermal gradient is too large, then the surface temperature will be too high (a minimum average temperature must be reached to give the desired amount of sintering or melting). Several negative effects are associated with an excessively high surface temperature. These effects include: 1) thermal degradation and/or oblation of the surface, 2) balling of the surface, and 3) excessive heat loss from the surface due to radiation and convection.2.5 Controlling the Thermal GradientIn early SLS experiments, it was observed that the thermal gradient was too high and steps were necessary to reduce the gradients into the acceptable range.

The excessive thermal gradient was manifested by a balling phenomenon (at that time called the "spider web structure") and by degradation of the material (2). To relieve these problems, a combination of the approaches presented below was used.The issue of controlling thermal gradients within the layer must be addressed as we work to increase the scan rate (area scanned per unit time). An increase in the scan rate requires an increase in the rate of laser energy input to the part. To do this, it is desirable to use more laser power and to reduce the proportion of the laser energy that is lost from the surface due to oblation, radiation and convection. A better understanding of the factors affecting the thermal gradients and methods of limiting the thermal gradients is necessary.Several approaches can be used to reduce the thermal gradient:1)Increase the part bed temperature.

This reduces the laser beam energy that is required to accomplish the sintering (or melting) of the powder and therefore the thermal gradient. The part bed temperature is limited due to part growth or caking of the part bed (non selective sintering of the powder outside of the part).153(3)2) Reduce the extinction coefficient. In many cases we do not have control over the optical properties of the powders. Also, this strategy is limited by the increase in reflectivity of the part bed as the extinction coefficient becomes small with respect to the particle size.3)Increase the laser beam interaction time.

This can be done by a) reducing the linear scan speed and reducing the laser power, b) increasing the laser beam spot size, or c) using more beam overlap (less spacing between the scan lines for a given laser beam spot size).Once a material is selected and the maximum part bed temperature has been determined, then increasing the laser beam interaction time is the primary tool that can be used to reduce the thermal gradient. If increasing area scanned per unit time is an objective, then a reduction in the linear scan speed is not an attractive option. Therefore the two approaches that appear most attractive are increasing the laser beam spot size and using more beam overlap.2.5.1Increasing Beam OverlapCurrently in SLS, it is common to use a beam overlap of 2 to 4 (the spacing between adjacent scan lines is 1/2 to 1/4 of the laser beam spot size). Each point to be sintered (or melted) is hit by the beam 2 to 4 times.

Each point to be sintered (or melted) is hit by the beam 2 to 4 times. Heat is conducted away from the surface into the layer between passes of the beam and this reduces the thermal gradient. Heat is also radiated and convected away form the surface during this time. For these reasons, the time between scan lines is a significant factor in determining the thermal gradient and the degree of sintering (or melting) of the layer. Figure 4 shows the temperature at a point as a function of time and the influence of time between scans and the maximum temperature. If a simple parallel line fill pattern is used, then the time between scans may not be uniform across the crossection (Fig. 5). Differences in material properties have been observed due to variation in the time between scans. For these reasons, it is desirable to take steps to bound the time between scans.Current commercial SLS machines use a spot size of 0.4 mm.

For these reasons, it is desirable to take steps to bound the time between scans.Current commercial SLS machines use a spot size of 0.4 mm. Increasing this spot size would increase the laser beam interaction time for a given scan speed, and the scan spacing could be increased while maintaining the same scan overlap. The disadvantage of this approach is that it increases the minimum feature size. For this reason, a dual beam size system is desirable. The small beam would be used for outlining and small features; and the large beam would be used to fill larger areas.2.5.2 Increasing the Laser Beam Spot Size3. OBJECTIVEThe main objective of the SLS research at Clemson University is to further develop an understanding of the SLS process with an emphasis on the energy delivery system. Of particular interest is the effect of scanning parameters and techniques on strength, speed and system robustness.

Of particular interest is the effect of scanning parameters and techniques on strength, speed and system robustness. This understanding will be used to speed up the scanning process while producing parts with equal strength and accuracy.Limiting the thermal gradient in the layer during the laser material interaction is a key issue in SLS. Addressing this issue will lead to an increase in the scan rate (3).3.1 Hypothesis3.2 ApproachAn experimental approach is being used. Test bars are being fabricatedand strength tested. Designed experiments and parameter trade-off studies are being conducted.An SLS process research workstation has been built. This machine is similar to commercial SLS machines in its general configuration.

This machine is similar to commercial SLS machines in its general configuration. The primary differences are: 1) this machine incorporates the unique feature of a variable beam spot size, 2) the oven has been enlarged horizontally in an effort to reduce the effect of the oven walls and the feed heaters on the part temperature distribution, and 3) the oven height has been reduced to reduce natural convection.3.2.1 Equipment1541.2,3.isFigureFigureScan patternPartially filled outline showing beam overlap156Temp.TemperatureDepth into bedBed temp.Figure 3. Temperature vs. depthTimeFigure 4. Time vs. temperature at a point157IA ~ IA~ 14~ 14~ 14~~ ~ I~~~,~,~Ir ,, ,If~• ,IrFigure 5. Variation in vector length as a function of geometry. This leads to a variation in the time between scans.158.

PROVED MICROSTRUCTURE AND INCREASED MECHANICAL PROPERTIES OF ADDITIVE MANUFACTURE PRODUCED TI-6AL-4V BY INTERPASS COLD ROLLING.Filomeno Martina∗,a, Stewart W. Williams∗, Paul Colegrove∗∗Welding Engineering and Laser Processing Centre (WELPC), Cranﬁeld University, Cranﬁeld, MK43 0AL, United Kingdom af.martina@cranﬁeld.ac.ukAbstractDistortion, residual stress and mechanical property anisotropy are current challenges in ad- ditive manufacturing (AM) of Ti–6Al–4V. High-pressure, interpass rolling was applied to linear AM parts and resulted in a change from large columnar prior β grains to a completely equiaxed microstructure with grains as small as 89 µm. Moreover, α laths thickness was also reduced to 0.62 µm. The change in material microstructure resulted in a substantial improvement of all me- chanical properties tested, which were also totally isotropic.

The change in material microstructure resulted in a substantial improvement of all me- chanical properties tested, which were also totally isotropic. In rolled specimens, maximum mea- sured strength and elongation were 1078 MPa and 14% respectively, both superior to the wrought material. Distortion was reduced to less than half. Rolling proved to be a relatively easy method to overcome some of the critical issues which keep AM from full industrial implementation.1 IntroductionThe anisotropy of mechanical properties is one of the issues currently affecting additive man- ufacturing (AM) [1, 2]. Due to solidiﬁcation characteristics of Ti–6Al–4V, parts microstructure is characterised by large columnar prior β grains, parallel to the thermal gradient, hence orientated along the vertical direction.

These grains are also strongly textured [3].Residual stress and distortion are also present challenges for AM [4, 5]; at the moment, they are overcome with post deposition heat treatments [6]. Heat treatments also inﬂuence the microstruc- ture (consequently the mechanical properties) and result in an increase of the elongation without compromising the ultimate tensile strength [7]. However, anisotropy is not addressed.In welding research, many techniques have been investigated to mitigate residual stress and distortion, including tensioning [8], peening [9, 10] and rolling [11, 12]. The latter can be imple- mented easily and was attempted on steel AM structures for the ﬁrst time by Colegrove et al. [4], resulting in a substantial reduction in distortion and residual stress, as well as grain reﬁnement.

[4], resulting in a substantial reduction in distortion and residual stress, as well as grain reﬁnement. Hence the same technique has been applied on Ti–6Al–4V deposits to investigate whether similar results could be achieved.2 Experimental methodsA standard TIG power source (Lincoln Electric Invertec V310–T) was ﬁtted onto a in-house designed rolling rig, as shown in Fig. 1. This deposition process belongs to wire+arc additive manufacturing (WAAM), which in general consists in the use of arc welding tools for additive manufacturing purposes. Linear structures were deposited onto a 6 mm thick substrate, using the parameters listed in Table 1. Rolling started and ﬁnished 35 mm within the deposit, and was done between passes, with a speed of 3 mm s−1 after the material had cooled to room temperature. Three loads were applied: 50 kN, 75 kN and 100 kN.

Three loads were applied: 50 kN, 75 kN and 100 kN. Walls built for microstructural analysis, as well as measuring out-of-plane distortion and geometric characteristics, were 370 mm long. Parts were sectioned at three locations to increase measurement reliability. The same cross sections were used for microhardness measurements (every 0.5 mm starting from 2.5 mm into the base plate with a load of 500 g and indentation time of 15 s).

The same cross sections were used for microhardness measurements (every 0.5 mm starting from 2.5 mm into the base plate with a load of 500 g and indentation time of 15 s). Three additional walls were built for tensile testing [13]; dog bone specimens were extracted in both vertical and horizontal directions and had a length of 120 mm, with a gauge of 24 mm × 6 mm × 4 mm.Figure 1: Experimental setup for rolling investigation.Table 1: Deposition parametersWire feed speed Travel speed Peak current Background current Average current Pulse duration Frequency1.6 m min−1 4.5 mm s−1 150 A 70 A 110 A 0.05 s 10 Hz3 Experimental results and discussion3.1 Distortion and geometryDistortion reduced with increased load: the maximum out-of-plane distortion of the control specimen was 7 mm, and it decreased to 5 mm with a load of 50 kN, and to 3 mm with a load of 75 kN.

The plastic deformation induced by rolling resulted in changes to the specimens’ width and height, the former increasing with the load, the latter showing an opposite behaviour (Table 2). It’s worth noting the reduction in the standard deviation of layer height, which would aid industrial implementation of AM as it reduces variability.Table 2: Changes in wall width and layer height.Wall widthLayer heightControl 50 kN 75 kNAverage 5.71 mm 6.17 mm 6.71 mmStd. dev. Average 1.13 mm 1.04 mm 0.93 mm0.18 0.11 0.14Std. dev. 0.19 0.12 0.093.2 MicrostructureAll specimens are characterised by a Widm¨anstatten microstructure, however the prior β grains conﬁguration differs substantially between unrolled and rolled samples (Fig. 2). The former showed the columnar grains conﬁguration that is typical of AM deposits, with elongated grains a few millimetre wide and a length that often traverses the whole component (Fig. 2a).

2). The former showed the columnar grains conﬁguration that is typical of AM deposits, with elongated grains a few millimetre wide and a length that often traverses the whole component (Fig. 2a). The latter, instead, were characterised by a reﬁned microstructure with equiaxed prior β grains. In fact, when rolling at 50 kN the average grain size was 125 µm (Fig. 2b), and for the 75 kN specimen it was 89 µm (Fig. 2c). It must be said that even the rolled specimens had a region with columnar prior β grains, speciﬁcally the top ca. 2 mm. This region was also rich in martensite, in all specimens, rather than Widm¨anstatten which is observed below. The horizontal bands that corresponded to the deposition of each layer were observed in all specimens. For an explanation of the origin of these bands, please refer to Martina et al. [14].In terms of α phase, rolled specimens were still characterised by the repetitive pattern in α lamellae thickness, within each horizontal band [14, 15].

[14].In terms of α phase, rolled specimens were still characterised by the repetitive pattern in α lamellae thickness, within each horizontal band [14, 15]. In fact, α lamellae thickness was min- imum around the bottom of each band (i.e. near the base plate), and increased gradually until it reached its maximum just before the next band started (i.e. towards the top of the wall). However, rolling resulted in a reduced α lamellae average thickness: for the control specimen it was 0.79 µm, for a rolling load of 50 kN it was 0.66 µm, and for the 75 kN specimen it was down to 0.62 µm. The nucleation of α phase lamellae starts at grain boundaries; as the part cools down, lamellae grow until they hit each other, and then new lamellae start growing perpendicular to the already developed ones, until an equilibrium fraction of α phase is reached [16, pp. 32–33].

32–33]. Given the reﬁned prior β grains, it is possible that the distance covered by α lamellae before they hit each other was shorter, hence their overall size was reduced.(a)(b)(c)Figure 2: Optical microscopy images of top areas from (a) control specimen, (b) specimen rolled at 50 kN and (c) specimen rolled at 75 kN.3.3 Mechanical propertiesRolling resulted in an improvement of all mechanical properties tested, which were always higher for both horizontal and vertical directions, apart from the elongation of the vertical unrolled specimens. In Fig. 3 tensile test results from the present research are compared against the wrought material tested by Wang et al. [2], and the speciﬁcations minima prescribed by ASTM B265 – 09a [17]. For the specimen rolled at 75 kN all properties were better than the wrought material.

[2], and the speciﬁcations minima prescribed by ASTM B265 – 09a [17]. For the specimen rolled at 75 kN all properties were better than the wrought material. The hardness values were 374 HV and 377 HV, for the 50 kN and 75 kN specimens, respectively; both values are higher than the unrolled sample (367 HV).The improvement in strength and hardness could be due to two reasons. First, the reﬁnement in the α phase is associated with an improvement of these two mechanical properties, at least for the cooling rates experienced in the WAAM process [18]. Second, each rolling pass induces a certain amount of cold work. Even if the subsequent deposition effectively has an annealing effect, this could be only partial. In fact, when rolling at 100 kN was attempted, the part fractured during rolling of the seventh layer.As for the elongation, a third factor must be considered: the reﬁnement of prior β grains, which has a detrimental effect of this property (so does work hardening).

On the contrary, the reduction in α lamellae thickness is beneﬁcial [18], and from case to case one of these factors would prevail over the others. For instance, in the case of the horizontal direction, elongation ﬁrstly decreases (from control to 50 kN) then increases (from 50 kN to 75 kN). Clearly for the 50 kN specimen, the contribution of cold work and reduced prior β grain size is dominant; however, for the 75 kN specimen the reduction in α lamellae thickness must be the factor with the greater effect, given that the elongation increases, even if only slightly. In the vertical direction the reﬁnement of priorYield strengthSpec.

In the vertical direction the reﬁnement of priorYield strengthSpec. minima ASTM B265-09Wrought (bars)WAAM control (horizontal)WAAM rolled at 50 kN (horizontal)WAAM rolled at 75 kN (horizontal)WAAM control (vertical)WAAM rolled at 50 kN (vertical)WAAM rolled at 75 kN (vertical)7007508008509009501000105011000510152025Yield / Ultimate tensile strength (MPa)Elongation (%)Ultimate tensile strengthElongationFigure 3: Comparison of rolled WAAM specimens against unrolled ones (error bars indicate stan- dard deviation), speciﬁcation minima [17], and wrought Ti–6Al–4V [2].β grains gives the dominant contribution: a much larger reduction in grain size is observed along this direction, and elongation always decreases with the load.Finally it is worth mentioning that cold work is the least important of the causes. If this wasn’t the case, a larger reduction in elongation would be seen in the 75 kN specimen, which had the highest strain.

If this wasn’t the case, a larger reduction in elongation would be seen in the 75 kN specimen, which had the highest strain. However elongation decreases marginally in the vertical direction and surprisingly increases in the horizontal one, which suggests that microstructural changes dominate over cold work.The most important ﬁnding of the present research is possibly the achievement of a totally isotropic mechanical behaviour.

In fact, already for a load of 50 kN the difference between the properties in the two tested direction is minimal; this is because even if the 50 kN specimen was al- ready characterised by an equiaxed prior β grains structure, possibly its texture wasn’t completely randomised [19] — differently from the 75 kN one, whose behaviour is completely isotropic.4 ConclusionsRolling of Ti–6Al–4V wire+arc additively manufactured walls resulted in:reduced distortion, because of changes in the geometry of deposits and most likely reduction in residual stress;reduced variability of layer height;reﬁned prior β grains and α lamellae, and change from columnar to equiaxed microstructure due to recrystallisation of the β phase;improved mechanical properties — strength, hardness and elongation were almost always better in the rolled case, and properties of the 75 kN specimen were even better than the wrought material ones;isotropic mechanical properties.AcknowledgementsThe authors would like to thank Mr Brian Brooks and Mr Flemming Nielsen for their support during the research activities.

proved Quality of 3D-Printed Tissue Constructs Through Enhanced Mixing of Alginate HydrogelsDaniel L. Cohen,1Andrew M. Tsavaris,1,2 Winifred M. Lo,1,3 Lawrence J. Bonassar,1 Hod Lipson11Cornell University, Ithaca, NY, 2Princeton University, Princeton, NJ 3 Northwestern University, Evanston, ILABSTRACTWhile alginate hydrogel is a desirable material platform for Solid Freeform Fabrication (SFF) of cell-seeded tissue engineering scaffolds, achieving consistently high-quality results can be challenging. Local variations in the material properties cause inconsistent material deposition behavior and consequently decrease the resultant geometric fidelity of the construct. The effects of gel mixing on material property consistency, geometric fidelity, and cell viability were characterized in an attempt to improve the formulation’s compatibility with SFF processing.

The effects of gel mixing on material property consistency, geometric fidelity, and cell viability were characterized in an attempt to improve the formulation’s compatibility with SFF processing. Material homogeneity was quantified through a novel experimental setup composed of an EnduraTEC mechanical test-frame and custom syringe-extrusion jig. Cell viability and geometric fidelity were assessed using standard protocol. The baseline mechanical stiffness of the printed samples was 16±3 kPa (n=6).

Cell viability and geometric fidelity were assessed using standard protocol. The baseline mechanical stiffness of the printed samples was 16±3 kPa (n=6). We found that increasing mixing reduced material inconsistency and improved geometric fidelity, without adversely affecting cell viability: the printed construct quality was drastically improved by increasing mixing well beyond previously established limits.IntroductionSFF is a powerful tissue engineering technique; compared with scaffolds produced through traditional tissue engineering techniques, SFF delivers scaffolds with increased geometric complexity, spatial heterogeneity, and patient-specificity. Many previous SFF efforts have been limited by high variations in print quality due to material property inconsistency.1 To address this limitation, it is necessary to develop a platform for quantifying heterogeneity of SFF inks.

As such, the objectives of our study were to (a) develop a framework for quantifying heterogeneity of the material properties of SFF inks by measuring temporal variation in load during the printing process, (b) determine the effect of mixing intensity on variations of mechanical properties and geometric fidelity of printed parts, and (c) verify maintenance of cell viability.MethodsAlginate Gel PreparationGels were created by mixing a 2:1 ratio of 2% alginate and 1% calcium sulfate (a cross- linking agent), both in PBS, by pumping the reagents back and forth through a three-way stopcock(Fig. 1). Gels were loaded into a syringe and allowed to cure for 30 minutes between the completion of mixing and the beginning of testing. Four different curing times were characterized: 15, 30 45 and 60 minutes.

Four different curing times were characterized: 15, 30 45 and 60 minutes. Thirty minutes was the minimum curing time that resulted in a gel with sufficiently high stiffness for SFF; curing times less than thirty minutes resulted in gels which did not hold their shapes upon deposition.In some mixtures, the gel at the nozzle of the syringe had solidified more than the gel present in the barrel of the syringe; this first bit of gel was expelled from all syringes prior to starting all tests. This was performed because it was observed that the viscosity and the resistance exhibited by this particular portion of the gel was not representative of the majority of the material in the barrel of the syringe. A potential explanation for this phenomenon is that the gel closest to the nozzle was the portion most exposed to air and evaporation occurred.

A potential explanation for this phenomenon is that the gel closest to the nozzle was the portion most exposed to air and evaporation occurred. Thus, with less water present in the gel composition, the material extruded might have exhibited more solid-like behavior.Mechanical Noise TestingGel was loaded into a syringe and mounted on an EnduraTEC mechanical test-frame (ELF 3200; Bose Corporation; Eden Prairie, MN) using a custom syringe-extrusion jig (Fig. 2). In each test, gel was extruded by depressing the syringe piston at a constant rate of 0.25 mm/s, resulting in a volumetric flow rate of about 0.043 mL/s. Throughout each test, the piston of the syringe was moved a total of 9.5 mm. Extrusion occurred through a tapered syringe tip with a diameter of 0.010 in. The resultant load, measured to within 0.01 N by a 50 lb load cell, was recorded at a frequency of approximately 200 Hz.

Extrusion occurred through a tapered syringe tip with a diameter of 0.010 in. The resultant load, measured to within 0.01 N by a 50 lb load cell, was recorded at a frequency of approximately 200 Hz. Tests were run on gels which were mixed to varying extents, as defined by the number of times the gels were pushed through the stopcock: 8, 16, 32, 64, 128, and 200 times.Fig. 1 Mixing process used to create alginate gel with two syringes and a three-way stopcock. The solutions were pumped back-and-forth through the stopcock. Each stroke counted as one time that the gel was mixed; a back-and-forth cycle would count as two mixings.The output extrusion force vs. time data was run through a band-pass filter to eliminate irrelevant low and high-frequency noise.

time data was run through a band-pass filter to eliminate irrelevant low and high-frequency noise. The high frequency noise was attributable to measurement and environmental noise, whereas the low frequency component was resulting from the one degree internal tapering of the syringe barrel which is a manufacturing artifact of the syringe.The standard deviation of the linear region of this force vs. time data was named “mechanical noise.” In order to characterize the sample-to-sample variation in the levels of mechanical noise, the standard deviation of the mechanical noise was measured.Fig. 2 Custom syringe-extrusion jig used in mechanical testing.Cell Viability Testingthe condyles and tests, articular chondrocytes were harvested from patellofemoral groove of bovine joints collected 1-3 days prior. Cells were isolated via digestion using a 0.25% collagenase solution in media over 18 hours and seeded in a 2% alginate gel in PBS the following day.

Cells were isolated via digestion using a 0.25% collagenase solution in media over 18 hours and seeded in a 2% alginate gel in PBS the following day. Gels were cross-linked in a 2:1 volume ratio with 1% CaSO4, allowed to cure for periods of time ranging from 15-45 minutes, and then extruded through a syringe at a rate of 0.024 mL/s. The effect of tip diameters was also examined by extruding the gel through syringe tips including 0.010 inch, 0.023 inch, and 0.200 inch diameters. Printed samples were incubated in media overnight at 37º C. Cell viability was assessed using fluorescent microscopy and the LIVE/DEAD cell viability assay the following day.For viabilityRapid PrototypingFor rapid prototyping, gels were mixed and transferred to syringes compatible with the Fab@Home rapid prototyper (Fig. 3). Prior to printing, gels were allowed to cure for periods of time between 15 and 60 minutes.

3). Prior to printing, gels were allowed to cure for periods of time between 15 and 60 minutes. Desired geometries for printing were uploaded to the Fab@Home software in the form of STL files. Gels were then deposited along prescribed paths using the Fab@Home machine.Fig. 3 (left) The Fab@Home rapid prototyper; (right) the custom deposition head built for use of sterile syringes.ResultsMechanical StiffnessThe baseline mechanical properties of the printed gels, after the standard post-crosslinking treatment, was 16±3 kPa (n=6). This stiffness was determined using a standard confined compression test on the EnduraTEC mechanical test frame.0.60.50.40.30.20.1y = 0.0117x - 0.002 R2 = 0.9645Fig. 4 Representative plot of stress-strain curve generated by EnduraTEC stress testing. The linearly-fit function is also displayed along with the correlation coefficient.Mechanical NoiseLoad vs.

4 Representative plot of stress-strain curve generated by EnduraTEC stress testing. The linearly-fit function is also displayed along with the correlation coefficient.Mechanical NoiseLoad vs. time data generated during the extrusion tests on the EnduraTEC mechanical test frame reflected material heterogeneity in addition to added noise from the machine and the syringe’s changing inner diameter (Fig. 5, left). For example, while material heterogeneity was discernible between tests, a downward sloping trend reflecting a gradual increase in the amount of load required to maintain the same volumetric flow of material was observed. This trend was likely resultant from the one degree internal slope of the syringe barrel, i.e. manufacturing-related draft. The draft obscured the component of the data which resulted from material heterogeneity; consequently, quantifying noise via standard deviation of an average load cell response would not produce meaningful results.

As such, it was decided that a band-pass filter, eliminating both high-frequency noise from the load cell and low-frequency noise from the syringe, was necessary to process the raw data prior to further analysis.Load Cell Readings Over Time (200X)Band-pass Filtered Load Cell Readings)N( g n d a e Ril le C d a o Ll le C d a o L d e r e t l iF)N(s g n d a e Ri2000 4000 6000 8000 Time Point (1/200 sec)2000 4000 Time Point (1/200 sec)60008000Fig. 5 (left) Representative raw load vs. time data for gel mixed 200 times; (right) Representative band- pass filtered data for gels mixed 200 times [green] and 8 times [black].00.0010.0020.0030.0040.0050.0060.0070.0080.0090While it is visually apparent that there is less temporal variation in the mechanical properties of gels that were mixed more as evidenced by narrower load reading bands (Fig.

5, right), this “mechanical noise” was quantitatively defined as the standard deviation of the load values over time (Fig. 6).)N( e s o Nila c n a h c e Mi10.80.60.40.200100# of Mixings200)N( e s o Ni. h c e Mf o .D S.0.60.50.40.30.20.100100 # of MixingsFig. 6 Effect of mixing intensity on gel homogeneity. (left) “Mechanical noise,” i.e. standard deviation of temporal variation in load necessary to print; (right) Sample-to-sample variation in mechanical properties of gel, i.e. standard deviation of mechanical noise.Gels mixed to a lesser extent (i.e. 8 times) exhibited substantial mechanical noise, whereas those mixed to a greater extent (i.e. 200 times) exhibited significantly less mechanical noise. Moreover, sample-to-sample variation decreased significantly as the mixing duration increased.

200 times) exhibited significantly less mechanical noise. Moreover, sample-to-sample variation decreased significantly as the mixing duration increased. This was judged from the standard deviation of the mechanical noise experienced for each treatment.Cell ViabilityIncreased magnitudes of mixing did not appear to exert a significant impact on cell viability in the seeded gels. Indeed, across magnitudes of 8, 64, and 200 times of mixings, cell viability appeared to remain relatively constant, at a magnitude of 74% (± 10-15% S.E.M.) when viability was normalized to the initial cell viability (Fig. 8). Moreover, decreasing the extrusion tip diameter from 0.200 inch to 0.010 inch also appeared to have negligible impact on cell viability. Finally, increasing the gel curing time across magnitudes of 15, 30, and 45 minutes appeared to have insignificant impact on cell viability.200Fig. 7 Representative images of fluorescent microscopy using LIVE/DEAD viability assay.

7 Representative images of fluorescent microscopy using LIVE/DEAD viability assay. (left) Live; (right) Dead.While the number of mixings, the printing tip diameter, and the curing time all did not negatively impact cell viability, it appeared that increasing mixing may actually improve cell viability in special circumstances. Gels mixed 200 times exhibited higher levels of cell viability than those mixed 8 times, when the gels were allowed to cure for a period of 45 minutes. One explanation for this phenomenon is that higher levels of mixing facilitate a better distribution of the cross-linking agent, calcium ions, throughout the alginate gel and allow for more homogeneous solidification during the 45-minute curing period. Thus, better-mixed gels experience lower peak- shear during extrusion and have a lesser tendency to kill cells throughout the printing process.Fig.

Thus, better-mixed gels experience lower peak- shear during extrusion and have a lesser tendency to kill cells throughout the printing process.Fig. 8 Effect of (left) tip diameter and (right) curing time on cell viability as measured 24 hours after printing.Geometric FidelityGels mixed for greater duration had a tendency to deposit at a constant rate, rather than agglomerating at the extrusion tip before falling uncontrolled onto the print surface; this continuous flow resulted in complete parts without material voids (Fig. 9, top-right).Fig. 9 (top-left) A sample print of a gel mixed 32 times exhibiting discontinuous material deposition streams and a resultant incomplete geometry with material voids (top-right) Complete sample print of a gel mixed 200 times.

(bottom-left) CAD model of ovine meniscus, (bottom-center) printed ovine meniscus with 32X mixed gel, and (bottom-right) printed ovine meniscus with 200X mixed gel.Not only was deposition constant, but the texture of the printed material was more uniform for those gels mixed to greater extent. Finally, gels printed after increased mixing generated shapes closer to the desired geometry, i.e. had greater geometric fidelity (Fig. 10).Fig. 10 Sample, high-fidelity print of a space shuttle (1” long).DiscussionThe goal of this research was to identify a platform for quantitatively evaluating the heterogeneity of printing inks and to facilitate the translation of alginate scaffold construction from the domain of injection molding to that of SFF.

The latter goal was facilitated by studying the impact of mixing on gel homogeneity, as well as on the retained viability of cells seeded throughout the scaffold.Analyzing the load cell response during extrusion of alginate gels appeared to generate a good measure of the material’s heterogeneity. The EnduraTEC was highly sensitive to the mechanical noise of the material, as well as interference from the syringes used during the test and the machine itself. These irrelevant components were removed using band-pass filtering of the raw data, while still retaining information about the material response. When processing the data, the option of subtracting baseline data—obtained from trials run using the same syringes filled with atmospheric air or water—was also considered. However, attempts to do so resulted in processed data which still exhibited the downward trend incurred during the test as a result of the syringe’s construction.

However, attempts to do so resulted in processed data which still exhibited the downward trend incurred during the test as a result of the syringe’s construction. One possible reason for why subtracting baseline data may have been an inadequate means of dealing with the trend of increasing force magnitudes is the fact that the starting position of the syringe varied from test to test. Additionally, the resistance experienced by each syringe varied depending upon the number of mixings the barrel had previously experienced. Thus, it was decided that using a band-pass filter to eliminate the high-frequency noise (resultant from the machine’s interference) and low-frequency noise (resultant from the syringe draft) was necessary.Each syringe contained sufficient gel (6 mL) to run two separate tests. These two tests were run immediately after one another. There was no apparent pattern in behavior separating the gel’s performance between the first and second tests.

These two tests were run immediately after one another. There was no apparent pattern in behavior separating the gel’s performance between the first and second tests. Consequently, to analyze the behavior from each syringe, we averaged the mechanical noise values from both tests and took the average to be the representative value from that particular sample.Increasing the duration of mixing appeared to better distribute the cross-linking agent throughout the gel, thus accounting for the improved uniformity of the material texture and viscosity. Qualitatively, the gel appeared to be smoother and more uniform in color and texture. Additionally, the material printed more easily on the Fab@Home than gels mixed for a low level of mixings. The sputtering and unpredictable deposition patterns previously observed were mitigated by increasing the duration of mixing.

The sputtering and unpredictable deposition patterns previously observed were mitigated by increasing the duration of mixing. Finally, it appeared that gels which were mixed more experienced less phase-separation upon, and after, deposition.Results from mechanical testing indicate that less mechanical noise is experienced during the printing process for those gels mixed for greater duration. These gels also tend to require less force for extrusion. This suggests that those gels mixed for shorter duration and with less uniform distribution of cross-linking agent likely experience clogging at the extrusion tip. When clumps of gel solidify within the barrel of the depositing syringe, the solid clumps tend to be trapped in the syringe tip during the extrusion process.

When clumps of gel solidify within the barrel of the depositing syringe, the solid clumps tend to be trapped in the syringe tip during the extrusion process. This reduces the overall volume of alginate gel leaving the syringe for a given force, thus requiring the machine to exert greater force to extrude the gel at the same volumetric flow rate.Mixing did not appear to have a significant detrimental impact on cell viability, despite concerns from experts in the field that more mixing would cause the seeded cells to induce apoptosis.ConclusionsThis study demonstrated a novel framework for quantifying material property homogeneity of printing inks. Furthermore, increased mixing was found to drastically increase ink uniformity, and consequently, improve the geometric fidelity of printed parts while not affecting cell viability. By mixing the gel an order of magnitude more than previously accepted and published limits, the print quality was substantially improved.

By mixing the gel an order of magnitude more than previously accepted and published limits, the print quality was substantially improved. These resulting improvements clear the way for extended application of this tissue engineering technology with greater precision and reliability, while also introducing a novel method for optimizing other types of printing inks that present challenges related to material property inconsistency.References1. Cohen DL, Malone E, Lipson H, Bonassar LJ. Direct Freeform Fabrication of Seeded Hydrogels in Arbitrary Geometries . Tissue Engineering 12, 5, 2006.2. Cao, Y., Vacanti, J.P., Paige, K.T., Upton, J., and Vacanti, C.A. Transplantation of chondrocytes utilizing a polymer cell construct to produce tissue-engineered cartilage in the shape of a human ear. Plast. Reconstr. Surg. 100, 297, 1997.3. Malone E., Rasa K., Cohen D. L., Isaacson T., Lashley H., and Lipson H.

inhomogeneity in chondrocyte-seeded constructs. J. Orthop. Res. 23, 134, 2005.6. Malone E. and Lipson H. Fab@Home: The Personal Desktop Fabricator Kit, Rapid Prototyping Journal, 13, 4, 2007.AcknowledgementsThis research was supported by the Cornell Center for Materials Research, funding from the National Science Foundation (Grant DMR-0520404), the NSF Graduate Research Fellowship Program, and the James and Rebecca Morgan Tissue Engineering Initiative. The authors would also like to thank Satish Degala, Jeff Ballyns, and Tunde Babalola for their technical assistance throughout this project..

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceAn Improved Vat Photopolymerization Cure ModelDemonstrates Photobleaching EffectsMohammad Mahdi Emami1 and David W. Rosen 1,21 Digital Manufacturing and Design Centre (DManD), Singapore University of Technologyand Design, Singapore 487372, mahdi_emami@sutd.edu.sg2 George W. Woodruff School of Mech. Engineering, Georgia Institute of Technology, Atlanta, GA 30332, david.rosen@me.gatech.eduKeywords: Photopolymerization, Photobleaching, Grayscale Stereolithography, Volumetric IntensityAbstractAn improved high-fidelity simulation model for a grayscale projection micro-stereolithography process has been developed. The modeling purpose is to accurately predict cured part shapes and dimensions, given a radiation intensity distribution.

The modeling purpose is to accurately predict cured part shapes and dimensions, given a radiation intensity distribution. The model employs COMSOL to solve a series of chemical reaction differential equations that model the evolution of chemical species (photoinitiator, monomer, and polymer) concentrations. Additionally, the model incorporates the effects of oxygen inhibition and species diffusion. This research offers two primary contributions to the cure model: the consideration of volumetric intensity to model variations in photoinitiator absorbance as a function of depth into the resin and a change to the rate model for photoinitiator to free radical conversion. The effects of these changes demonstrate observed photobleaching effects. Simulated cured part profiles are compared to experiments and demonstrate good agreement. Additionally, initial results are presented on the usage of the simulation model in a new process planning method.1.

Additionally, initial results are presented on the usage of the simulation model in a new process planning method.1. IntroductionMonolayer photopolymerization with a grayscale mask can be used for fabrication of parts with a freeform surface profile in a single shot of exposure. Applications include micro-lens arrays and other optical components, as well as surface patterning with limited pattern height. Unlike layer- by-layer methods, the part is grown on a clear substrate by controlling the exposure energy distribution accordingly. However, obtaining such distribution that accurately cures desired profiles might be challenging and limits the accuracy regarding industrial application. Although the working curve is typically used in stereolithography processes, it only defines the relationship between exposure time and depth of cure [1,2]. However, for microfabrication of a freeform profile, lateral or side curing should be considered.

However, for microfabrication of a freeform profile, lateral or side curing should be considered. This is the reason that expensive photoresists with long process times (layer deposition, prebaking, and development) are preferred for microfabrication [3,4]. Shrinkage and oxygen inhibition are limited by utilizing photoresists in comparison to liquid photopolymers.Lateral curing or more precisely the prediction of a cured 3D profile is related to coupled factors and phenomena that occur during the irradiation process that makes modeling more complex. Some of these effects are oxygen inhibition, photo-bleaching, photo-trapping, self-focusing, and shrinkage [5]. All affected parameters and properties potentially can change the shape or dimensions of the final profile.1940In one of the first attempt to cover the oxygen inhibition effect, Jariwala et. al. built a semi- empirical model of photochemical reactions with consideration of oxygen inhibition [6].

al. built a semi- empirical model of photochemical reactions with consideration of oxygen inhibition [6]. Then, they used this model to propose a process planning method for generating a series of binary masks to create spherical and aspherical lenses [7]. The simulation result demonstrated an overestimation discrepancy between COMSOL simulations and the experimental working curve for curing depths of more than 60 m [6].Many researchers have studied and developed models for photobleaching (PB) effects. Depending on the nature of the photolysis products, the consumption of the photoinitiator can either lead to an increase in light intensity in the underlying layers (if the photolysis product is more transparent at the irradiating wavelengths) or a reduction in light intensity (if the photolysis product is strongly absorbing). Miller et. al. proposed a generalized model of PB for thick samples [8].

Miller et. al. proposed a generalized model of PB for thick samples [8]. They conclude that diffusion is expected to have negligible effects in most PB systems. However, their work does not include experimental evidence.This article builds on the foundation provided by researchers at Georgia Tech [6,7] to extend the photochemical model with a new differential equation for volumetric intensity to represent light intensity gradient as well as defining a new mechanism for initiator self-dissociation to incorporate the PB effect. Consequently, the proposed method can estimate the curing height for thick layers. The capability of this work has been demonstrated for the fabrication of microlens with sag heights of up to 800 µm.2. Chemical Kinematics of PhotopolymerizationFree radical photopolymerization includes the stages of initiation, propagation, and termination, and the phenomena of species diffusion, heat generation, and oxygen inhibition [9].

The process starts with initiation when the photoinitiator molecule dissociates under UV radiation to form initiator radicals. Subsequently, the generated radical reacts with monomer species, producing primary radicals.The process continues with the propagation step as monomer molecules react with the primary radical by a covalent bond and create an active polymer chain. The chain reaction continues until the termination step when radicals become extinguished or are otherwise unable to continue reacting. In the presence of dissolved oxygen, each active polymer chain or radical may react with oxygen to form a peroxide species, which converts them to an inactive polymer or dead radial. As a result, the propagation reaction is halted. The effect of oxygen inhibition described here is based on the works of Terrones-Pearlstein [10] and Dendukuri et. al. [11].In contrast, our new modeling scheme includes one more step called “self-dissociation” of initiator to cover the PB effect.

al. [11].In contrast, our new modeling scheme includes one more step called “self-dissociation” of initiator to cover the PB effect. Simulations without this step correctly estimate the curing depth only for layers less than 350µm thick. Beyond this, the estimation error increases by depth. The PB effect was distinguished by Jacobs as the major reason for the observed super-logarithmic behavior of some photopolymer systems [12]. Comparing the simulated results of cure depth with the working curve obtained experimentally shows an underestimation of curing depth (refer Section 6, Figure 5.a).In our photopolymer system (TMPTA and DMPA, see Section 5), the assumption is that the initiator is the only absorber of UV light. Also, a simple absorption test for pure TMPTA without initiator shows that it does not absorb UV light at 365nm. The initiators dissociate and consume1941during the curing and so their concentration gradually reduced.

The initiators dissociate and consume1941during the curing and so their concentration gradually reduced. So, the absorptivity of UV light changes over time of exposure and it allows more photons to penetrate and reduce the attenuation of light. This is called the photo-bleaching effect and is our primary hypothesis to explain the observed super-logarithmic behavior of our photopolymerSimulation results without self-dissociation indicated that the consumption rate of initiator is not high enough to demonstrate a PB effect. It means that other mechanisms might exist for converting the initiator to molecule without UV absorption. Photolysis of DMPA (initiator) showed that various chemical reactions and products were observed, depending on the intensity level [13]. Furthermore, dissociation may occur by the heat energy produced by polymerization. To model these effects, a new self-dissociation reaction equation is added to our COMSOL model.

Furthermore, dissociation may occur by the heat energy produced by polymerization. To model these effects, a new self-dissociation reaction equation is added to our COMSOL model. The abovementioned chemical reactions are summarized in Table 1.

Additionally, a self-dissociation rate of 𝐾𝑠𝑒𝑙𝑓 is defined that acts to reduce the initiator concentration; it has its greatest impact when the concentration of free radicals is large.To obtain a kinetic expression for the rate of polymerization, it is necessary to assume that 𝐾𝑝 and 𝐾𝑡 are independent of the size of the radical [14]. The rate of photochemical initiation 𝑅𝑖 is given by𝑅𝑖 = 2𝜑𝐼𝑣1where 𝜑 is referred to as the quantum yield for initiation and interpreted as the number of propagating chains initiated per light photon absorbed. The maximum value of 𝜑 is 1 for most photoinitiating systems. 𝐼𝑣 is the volumetric photon absorption rate, called the volumetric intensity in this paper (refer to section 4). The factor of 2 in Eq.

The maximum value of 𝜑 is 1 for most photoinitiating systems. 𝐼𝑣 is the volumetric photon absorption rate, called the volumetric intensity in this paper (refer to section 4). The factor of 2 in Eq. 1 is used to indicate that two radicals are produced per molecule undergoing photolysis.Finally, the rate of polymerization 𝑅𝑝 can be obtained from Eq.2 below [10,14]𝑅𝑝 = 𝐾𝑝[𝑀]√𝑅𝑖 2𝐾𝑡⁄2where [M] is the concentration of monomer and 𝐾𝑝 and 𝐾𝑡 are the propagation and termination rates, respectively. It is seen that Eq.2 has the significant conclusion of the dependence of the polymerization rate on the square root of the initiation rate. Doubling the rate of initiation does not double the polymerization rate; the polymerization rate is increased only by the factor √2.

Doubling the rate of initiation does not double the polymerization rate; the polymerization rate is increased only by the factor √2. This behavior is a consequence of the bimolecular termination reaction between radicals.To calculate the initial concentration of initiator [𝐼𝑛]0 and monomer double bond concentration [𝑀]0, the solvent density 𝜌 (𝑘𝑔/𝑚3) is required to convert the mass-based concentration (𝑘𝑔/𝑚3) to molar-based (𝑚𝑜𝑙/𝑚3). Considering a solution of n% (weight percent) of initiator where n is less than 5%, the volumetric change of solvent due to adding the initiator can be neglected. Hence, the total volume of solution can be considered as the total volume of solvent (Monomer). Therefore, the initial value of monomer concentration (double bond) is obtained[𝑀]0 = 𝑛𝐷𝐵ρ 𝑀𝑚3where ρ is the density of monomer, 𝑀𝑚 is the molar mass (molecular weight) of monomer, and 𝑛𝐷𝐵 is the number of double bond(s) in a monomer.

Note that based on acrylate type (mono or multi-functional) the number of double bonds should be considered in the above equation; for example, a tri-acrylate has three double bonds (𝑛𝐷𝐵 = 3). It shows that each monomer has three active components for polymerization.In the same way, the initial concentration of the Initiator is calculated as follows 𝑛 100 − 𝑛ρ 𝑀𝑖𝑛[𝐼𝑛]0 =4where n is the weight percent of initiator and 𝑀𝑖𝑛 is the molar mass of initiator.Finally, the simulation of the kinematic model provides a molar concentration of various chemical products in reactions. Consequently, the polymerization level, which is indicated by the degree of conversion (DoC), will be calculated at any arbitrary point simply by:1943𝐷𝑐𝑜𝑛𝑣 =[𝑀0] − [𝑀𝑑] [𝑀0]5Once the resin reaches to a certain amount of Dconv, it is converted to gel and the viscosity of the solution increases sharply, corresponding to a rapid transition from a liquid state to a solid state.

Therefore, Dconv provides a criterion for defining a boundary between liquid and solid states. Polymerization particularly in multifunctional monomer does not reach a complete conversion of reactants due to the sharp increase of viscosity as well as declining free volume. The ultimate Dconv can be measured by FTIR and is usually reported less than 0.7 in previous works. This restriction is also implemented in the proposed model. According to the knowledge of authors, limitation ofUsing the rate constants for TMPTA+DMPA listed in Table 3, a conversion degree value of 0.23 was determined by fitting to the data obtained experimentally. The value of 23 percent is consistent with the value measured by Boddapati via micro-rheology[15]. The parameters used in this study, as well as previous studies, are summarized in Table 3.3. Intensity Distribution ModelThe illumination of light on a surface can be represented by an intensity distribution model.

Intensity Distribution ModelThe illumination of light on a surface can be represented by an intensity distribution model. It will be used as an input of the volumetric intensity model which will be described in Section 4. The accuracy of the modeling depends on how well the predicted intensity profile matches the actual intensity. Two well-known approaches are available for this part – ray tracing method [16] and point spread function (SPF) [17]. The ray tracing approach is computationally intensive, while SPF may not incorporate the intrinsic non-uniformity of illumination in projection systems.In this research, the SPF method, also known as super-Gaussian, was chosen as a high- performance solution for small patterns. It has been implemented based on research published by Emami et. al. [17].

The concept of approximating the intensity of a pixel by a Gaussian distribution has been elaborated also by other researchers [18,19].Table 3- Estimated parameters of kinetic modeling for photopolymerization of TMPTAParameter (unit)Kp (m3/mol/s) Kt (m3/mol/s) KO (m3/mol/s)KSelf (m3/mol/s)𝜖wt%DoC%DescriptionPropagation rate Termination rate Oxygen inhibition rate Initiator Self-Dissociation rate Initiator molar absorption coefficient, decadic base Initiator Concentration Criterion for solidified resin Maximum intensityCurrent Researchwith PB 0.26 0.39 2.2w/o PB 0.26 0.39 1.81200013102%2%2323Boddapati[15]with O2 0.26 0.39 2w/o O2 1.61 14.34 -15155%5%2020Jariwala [7]1.66 1.31 125155, 20%20Imax (mW/cm2)18.218.214141.251944Figure 1- Gaussian Distribution of Intensity Profile, Ik(x).Based on the foregoing, the intensity distribution of a series of pixels on the image plane, 𝐼𝑠(𝑥),is given by Eq.6.𝑛 𝐼𝑠(𝑥) = ∑ 𝐼𝑘(𝑥) 𝑘=0𝑛= ∑ 𝐺𝑘𝐼𝑚𝑎𝑥𝑒−[4√2ln2 (𝑥−𝑘𝜌)/𝜔0]2𝑘=06where 𝐺𝑘 is the grayscale level of the kth pixel (integer value between 0 to 255), 𝐼𝑚𝑎𝑥 is the peak intensity of the full-on pixel, 𝜌 is pixel pitch, 𝜔0 is Gaussian half width, and n is the total number of pixels.

As shown in Figure 1, 𝐼𝑘(𝑥) falls rapidly when moving away from the Gaussian origin, line x = 0. To increase the computing performance of the superposition calculation, an effective area (EA) is defined as the area receiving 99.99% of the total exposure [17]. This area is bounded by x= ±2.15𝜔0.Figure 2- Normalized measured intensity profile vs. the super-Gaussian model for 20 pixels,Is(x).A parameter identification method has been used to find the model’s independent constants like as 𝐼𝑚𝑎𝑥, 𝜌 and 𝜔0. The light engine employed in this research is capable of providing a wide range of intensities: 1-100 (mW/cm2). 𝐼𝑚𝑎𝑥 is the average peak intensity of top-flat mask with Gray level of 255 (fully on mirrors) which is measured with a power meter.1945The remaining parameters (𝜌 and 𝜔0) were tuned so that the model output fits the intensity profile measured by a CCD camera. A series of linewidths were chosen to show the generality of the model.

A series of linewidths were chosen to show the generality of the model. The measured and calculated intensity distributions of a line pattern with a width of 100 pixels is shown in Figure 2. The intensity distribution of an irradiated mask composed of n pixels was calculated using Eq.6. Parameter optimization showed that the minimum discrepancy between the experiment and the model achieved when 𝜔0=18.8 µm and 𝜌 =21.6 µm.4. Volumetric IntensityIn physics, intensity is the power transferred per unit area, where the area is measured on the plane perpendicular to the direction of propagation of the energy. In photochemistry, any photon with sufficient energy can affect the chemical bonds of a chemical compound. Therefore, the intensity of absorbed light in moles of light quanta per volume per time, named volumetric intensity here, is required to calculate the number of affected bonds [10].In some previous works, the expression given for initiation rate Ri (see Eq.

1) does not exhibit dimensional consistency [5,9,15,20]. For unit consistency, it is required that 𝐼𝑣 has dimensions of moles per volume per time rather than an intensity (Einstein per unit area per time). The correct governing equations have been proposed in this section.ZLiquid Resin𝑧𝑑𝑧′𝑪𝒊𝒏(𝑥, 𝑧′, 𝑡)𝑧′𝑰𝒛(𝑥, 𝑧′, 𝑡)Volumetric Absorbed EnergyxX𝐼𝑠(𝑥)Figure 3- Volumetric Intensity Iv(x,z,t)The volumetric intensity 𝐼𝑣 is based on the Beer–Lambert law. Figure 3 shows a schematic of the curing process in which the light distribution at the surface is 𝐼𝑠(𝑥). The light penetrates the resin and the light intensity declines to 𝐼(𝑥, 𝑧́) based on the Beer-Lambert law. Derivation of 𝐼𝑠(𝑥) was described in Section 3. At a specific time, t, for any arbitrary infinitesimal layer 𝑑𝑧́, the concentration of absorber (initiator) is assumed to be constant in the z direction, or 𝐶𝑖𝑛(𝑥, 𝑧́, 𝑡) = 𝐶𝑖𝑛(𝑥, 𝑧́ + 𝑑𝑧́, 𝑡).

At a specific time, t, for any arbitrary infinitesimal layer 𝑑𝑧́, the concentration of absorber (initiator) is assumed to be constant in the z direction, or 𝐶𝑖𝑛(𝑥, 𝑧́, 𝑡) = 𝐶𝑖𝑛(𝑥, 𝑧́ + 𝑑𝑧́, 𝑡). In the other words, initiator concentration only varies along the x-axis.For this study, the initiator is assumed to be the only UV absorber. Furthermore, the effect of light refraction can be neglected since the UV light is projected perpendicularly to the substrate. Also, the resin is optically clear before and after curing. Therefore, the light scattering effect is negligible.1946In photopolymerization, the concentration of initiator 𝐶𝑖𝑛 is uniform at the start, but it is not uniformly changed during the process.

Therefore, the light scattering effect is negligible.1946In photopolymerization, the concentration of initiator 𝐶𝑖𝑛 is uniform at the start, but it is not uniformly changed during the process. For each wavelength, the local light intensity 𝐼𝑧 decreases along the beam direction 𝑧 according to the integrated form of Beer’s law [10,20]𝑧𝐼𝑧(𝑥, 𝑧, 𝑡) = 𝐼𝑠(𝑥, 𝑡) ∙ exp [− 𝛼 ∫ 𝐶𝑖𝑛(𝑥, 𝑧́, 𝑡)𝑑𝑧́]70 where 𝐼𝑠 is the incident light intensity at the outer surface of the reaction system and 𝐶𝑖𝑛(x,z,t) is the molar concentration of light-absorbing photoinitiator at the depth of 𝑧 at time t in the reaction system. The absorption coefficient of initiator 𝛼 varies with wavelength and temperature. Here, the temperature variation is not considered. Although the initial concentration distribution of the initiator is uniform through the reaction region it will not remaining steady and uniform during the reaction. The integral expression in Eq.

Although the initial concentration distribution of the initiator is uniform through the reaction region it will not remaining steady and uniform during the reaction. The integral expression in Eq. 7, corresponds to the variation of the initiator over the layer thickness. This spatial variation was not considered in some previous works or was averaged spatially [9,11,21]. Averaged concentration might be not accurate enough to demonstrate the PB effect.The amount of absorbed light intensity at the distance of 𝑧, 𝐼𝑎, is calculated as follows 8𝐼𝑎(𝑥, 𝑧, 𝑡) = 𝐼𝑠(𝑥, 𝑡) − 𝐼𝑧(𝑥, 𝑧, 𝑡)The variation of 𝐼𝑣 along the optical path direction, z, is obtained as the differential of 𝐼𝑎 withrespect to 𝑧 as follows𝐼𝑣 =∂𝐼𝑎 ∂z= −∂𝐼𝑧 ∂z= 𝛼𝐶𝑖𝑛(𝑧) ∙ 𝐼𝑧9To solve the above differential equation (Eq.9) and obtain 𝐼𝑧, the underneath boundary condition (BC) is considered. At the illuminated surface, and immediately above the substrate, the incident intensity is 𝐼𝑠(𝑥, 𝑡).

At the illuminated surface, and immediately above the substrate, the incident intensity is 𝐼𝑠(𝑥, 𝑡). We assume that at this surface neither reflection nor refraction is important. The BC at the incident light surface is defined by a Dirichlet BC, 𝐼(𝑥, 𝑧 = 0, 𝑡) = 𝐼𝑠(𝑥, 𝑡), which is a function of position (x) and time (t). Having a time-dependent intensity function adds the capability to simulate multiple masks during the illumination. The COMSOL ODE solver is employed to solve the volumetric intensity coupled to the rest of physics. Consequently, the attenuation of light changes during the illumination according to the concentration variation of photoinitiator.The molar absorptivity 𝜖, formerly called the extinction coefficient, is often used instead of 𝛼. It is related by 𝛼 = 𝜖 ∙ 𝑙𝑛(10) = 2.3𝜖 and result from the use of the Napierian instead of the decadic base, respectively, in the Beer-Lambert law. The value of 𝜖 used in this simulation was reported in Table 3.

The value of 𝜖 used in this simulation was reported in Table 3. The energy in a mole of light (Einstein) is calculated by 𝑁ℎ𝜈 or 𝑁ℎ𝑐/𝜆 where h is Planck’s constant, 𝑐 is the speed of light, 𝜆 is the wavelength and 𝑁 is Avogadro’s number [14]. Thus, the Eq.9 can be rewritten as follows𝐼𝑣 = 2.3𝜖 .𝜆 𝑁ℎ𝑐. 𝐶𝑖𝑛(𝑧) 𝐼𝑧10Here, the units of 𝐼𝑣 and 𝐼𝑧 are (𝑚𝑜𝑙 ∙ 𝑚−3 ∙ 𝑠−1) and (𝑊 𝑚2⁄) respectively. Thus, the differencebetween 𝐼𝑣 and 𝐼0 can be large for thick sections.19475. Experimental SetupTrimethylolpropane triacrylate (TMPTA), a trifunctional monomer with a technical grade containing 600 ppm of monomethyl ether hydroquinone (as an inhibitor), and 2,2-Dimethoxy-2- phenyl acetophenone (abbreviation - DMPA, trade name - Irgacure® 651) as a photoinitiator (Sigma-Aldrich) were used without further processing or purification. The weight ratio between the monomer and the photoinitiator was 98:2.

The weight ratio between the monomer and the photoinitiator was 98:2. The UV-curable resin was prepared by mixing these two compounds together and storing for one day until the initiator species was completely dissolved.In this research, a high-performance UV light engine (LE) from VISITECH was employed to illuminate high resolution and uniform grayscale patterns. The system specification is summarized in Table 4. This LE has an adjustable optical system designed that enables variable magnification with the capability of reducing or enlarging the projected image from the DMD. For the current research, a magnification ratio of 2:1 has been selected. Furthermore, the high frame rate of 19 KHz reduces the minimum achievable exposure time to the order of 100 µs.

For the current research, a magnification ratio of 2:1 has been selected. Furthermore, the high frame rate of 19 KHz reduces the minimum achievable exposure time to the order of 100 µs. In the other words, it increases the vertical resolution (in direction of z-axis) for higher intensities.Table 4- Specification of UV Light EngineSpecificationValueDMD resolution1920 × 1080 pixelsDMD typeMicromirror sizePixel size (Image Plane)0,95" 1080p UV Enhanced 10.8 × 10.8 (µm2) 21.6 (µm)Light sourceHigh Power LEDLight source wavelengthLight source powerContrast ratio365 (nm) 1 - 100 (mW/cm2) Typical 1300:1 (full ON/full OFF)Max binary frame rate19.5 (KHz)Figure 4- Experimental setup consists of a high-performance light engine, UV Camera, andPolymerization vat.1948The experimental setup is shown in Figure 4, along with a schematic of grayscale irradiation ofthe resin bath through the glass substrate.6.

Results and DiscussionTo illustrate the performance of the PB model, a comparison between the working curve with and without the PB effect is shown in Figure 5. The simulation result considering the PB effect can estimate the actual working curve of the resin up to a curing depth of 1100µm (Figure 5.b). This height provides adequate flexibility to design micro-optic parts. Without that, the simulation working curve only covers the curing depth less than 400µm. The experimental working curve of TMPTA shows super-logarithmic behavior. The proposed method with consideration of self- dissociation of initiator also shows the same behavior. However, the conventional method exhibits a logarithmic behavior.Figure 5- Comparison experimental working curve with simulation result without PB effect(blue dot line) and with PB effect (red dot line)Figure 6- Side image of a fabricated aspherical lens with a radius of 1620µm and sag height of 800µm.

Dash line and solid line represent the simulated profile with and without PB, respectively.1949To study the feasibility of the proposed model for thick grayscale monolayer fabrication, an intensity profile was chosen to fabricate spherical lens with a sag height of 800 µm. In parallel, this profile was imported into COMSOL and used as the input to predict the final cured profile with and without PB effect. Since the proposed model is an axisymmetric FEM simulation, only the cross-section of the fabricated lens is selected for comparison. The side image of the fabricated lens, as well as simulation results, are depicted in Figure 6.

The side image of the fabricated lens, as well as simulation results, are depicted in Figure 6. Although both simulated profiles demonstrate discrepancy, the error between the result with PB (dashed line) matches with the final experimental profile much better than the model without the PB effect.Figure 7- Simulation results of initiator concentration variation with PB (Kself =1200) and without PB (Kself =0)Figure 7 shows the consumption of initiator during UV exposure at the specific point of [100um,100um] in the simulated part shown in Figure 6 with and without the PB effect. A non- zero Kself value indicates that the PB effect is being modeled. Results demonstrate that the number of initiators consumed without the PB effect is very small compared to including the PB effect.7. ConclusionIn this work, an improved multiphysics model of photopolymerization consisting of chemical reactions and UV irradiation has been demonstrated.

ConclusionIn this work, an improved multiphysics model of photopolymerization consisting of chemical reactions and UV irradiation has been demonstrated. In the chemical reaction part, the photobleaching effect has been simulated by considering a new self-dissociation reaction. A comparison between the simulated and actual working curve demonstrated considerable improvement in estimating depth of cure up to 1100 µm. Furthermore, the proposed model capability was tested by fabrication of microlens with sag heights of 800 µm.Describing and utilizing a volumetric intensity instead of surface intensity provided a precise molar number of photons participating in the curing based on the surface intensity. It corrected the unit inconsistency shortcoming of the surface intensity that was observed in some previous works.Future work will utilize the improved photopolymerization model in a process planning methodto achieve high precision fabrication of optical components.19508.

provement in Geometrical Resolution of Plastic Laser Sintering by usingReduced Spot Sized LaserToshiki NIINO and Keisuke MORITAInstitute of Industrial Science, the University of Tokyo, Japan4-6-1 Komaba Meguro Tokyo, 153-8505 Japanniino@iis.u-tokyo.ac.jpABSTRACT Plastic laser sintering is one of the most promising processes for rapid manufacturingamong various additive manufacturing (AM) technologies. Though the process tends tobe applied to fabrication of larger parts in comparison to stereolithography, its ability ofcreating complex structure as an advantage of additive manufacturing technologiesshould be demonstrated in production of smaller parts and parts including fine and complex geometries as well. In this research, narrow CO2 laser beams with spot diameters of 130µm and 150µm were tested while the commercially available machinesare equipped with those around 500µm.

In this research, narrow CO2 laser beams with spot diameters of 130µm and 150µm were tested while the commercially available machinesare equipped with those around 500µm. Relationship between resolution (available wall thickness) and spot diameter is proportional when the diameter is greater than 150µm, but effect of reducing the spot size further is not significant. The minimum wallthickness of 180µm was obtained, but this part was so fragile that skill in breakouttreatment is critical. To discuss the mechanical strength of micro-plastic-laser-sintering,packing rate of obtained parts was introduced as an index of the strength. Buildparameter that minimizes the wall thickness without decreasing the tensile strengthbelow 30MPa was searched, and a set of parameters that provides minimal thickness of0.6mm was obtained. Reducing laser spot size inevitably leads to shrinkage of scanningrange of galvanometer mirrors.

Reducing laser spot size inevitably leads to shrinkage of scanningrange of galvanometer mirrors. To overcome this problem, the whole laser scanningsystem was set on an X-Y positioner which are driven by stepper motors. The wholeexposure area is divided into some regions each of which is smaller than range ofgalvanometer mirror system, and it is exposed by repeating fast scanning by galvanometer mirrors and slow sliding by the X-Y positioner. Problems occurring at the region boundary were investigated.

Problems occurring at the region boundary were investigated. As counter measures, overlapping of exposure areas and switching of region boundary are introduced and successfully eliminate the problem.NOMENCLATUREms: tm:mass of a sinterMelting temperaturetn: vb: : t:Nominal wall thickness Bulk volume of powder or a sinterPacking rate of powder or a sinterTrue density of powder materialINTRODUCTION Plastic laser sintering is one of the most promising rapid manufacturing methods amongvarious additive manufacturing technologies that have ever been developed andcommercialized. A wide choice of applicable thermoplastics facilitates fabrication offunctional parts or devices with various material performances such as heat resistance,mechanical strength and biodegradability[1]. On the other hand, in terms of geometricalpreciseness of fabricated parts, the process is inferior to some other additivemanufacturing processes such as stereolithography.

On the other hand, in terms of geometricalpreciseness of fabricated parts, the process is inferior to some other additivemanufacturing processes such as stereolithography. Because of this disadvantage,selective laser sintering tends to be used for fabrication of larger parts. Thus, the mostrapid manufacturing examples that have ever been developed are relatively large andincluding simple structures. However, additive manufacturing is more cost effectivewhen smaller parts are fabricated[2] since the cost is roughly proportional to processtime or volume of each part. Additionally, fabrication of more complex structure insmaller dimensions can give a better chance to demonstrate advantage of additivemanufacturing over the other conventional processes.The leading cause of the poor preciseness of plastic laser sintering is large laser beamdiameter in commercially available systems.

In a typical plastic laser sintering system, CO2 laser, whose wavelength is 10.6µm, is focused into a spot of 500µm in diameter while a typical stereolithography system employs near UV laser whose wavelength isaround 350nm and focuses into a small diameter of 50µm or smaller. Due to its longer wave length, to focus CO2 laser into a small spot is more difficult than UV laser, but there is still some room to reduce the diameter from theoretical point of view, and, in reality, we find a small focal spot less than 100μm in some CO2 laser applications such as laser drilling. In this research, the effect of reducing laser spot diameter onimprovement in preciseness of the plastic laser sintering as a process of machining isinvestigated. Two experimental setups equipped with optical systems that can focustheir lasers into small spots of 130μm and 150μm, respectively, were built and tested.Relationship between thickness of vertical wall and various process parameters areexamined.

Packing rate of the walls are measured as an index for their mechanical strength. MicroCT observation was performed to investigate distribution of pores in thethin walls.As mentioned previously, reducing spot diameter of CO2 laser to less than 100μm is not very difficult from the theoretical view point. However, it is not reasonable from theside of practice since it makes a sacrifice of scanning range of galvanometer mirrorsresulting in the shrinkage of the build envelope. To overcome this drawbgack, a hybridsystem consisting of galvanometers and stepper motors was tested (Fig. 1). In thissystem, the whole galvanometer mirror system including focusing optics is set on anX-Y positioner that is driven by the stepper motors. The stepper motors cannot scan thelaser spot as fast as the galvanometers but can provide greater range of movement moreeasily and economically.

The stepper motors cannot scan thelaser spot as fast as the galvanometers but can provide greater range of movement moreeasily and economically. Utilizing this advantage, we can expand the build envelope inscan and stitch way; we divide the whole exposure area into small regions, expose oneof them by high speed scanning with galvanometer mirrors, and convey thegalvanometer mirror system with the X-Y positioner to expose another region. Sincemoving the galvanometer system to the next resion takes time, there is a time lagbetween exposures of two neighboring resions. This difference may causes problemswhen a cross section which lays across a region boundary. This paper reveals theproblem and introduces its counter measure.Galvanometer mirror scanners and focusing opticsScanning range of a set of galvanometer mirrorsExpanded build envelopeExpanded build envelopeFig.

1 Schematic view for build envelope expansion byscan-and-stitchLASER OPTICS AND POWDER In this research, two sets of laser optics were used. One is a commercially available laser marker (ML-G9320, KEYENCE). This device, “scanner A” in the following description, is equipped with CO2 laser, and its beam is focused on the powder surface in a diameter of 150µm. Distance between lens and the focal plane is 113mm, andscanning range is 55mm x 55mm.The other one, “scanner B” in the following description, was designed specially for our purpose. This optics is also installed with CO2 laser. The beam emitted from the laser tube is once expanded to 17mm in diameter, then bent by galvanometer mirrors up to19° at the maximum and focused on the powder surface by a lens with focal length ofTbl.

3 Relationships between measured thickness and nominal thickness109mm. By this optic system, a 1/e2 diameter of 130µm is obtained. Parameters of laser optics employed in this research are summarized in Tbl. I. As the powder, ASPEX-PA produced by ASPECT Inc. was employed in the experiments of this research. This poweder is PA12 and has the average grain size of 45µm.FABRICATION OF FINE STRUCTURE BY REDUCED LASER SPOTRELATIONSHIP BETWEEN WALL THICKNESS AND SPOT DIAMETER Minimum thickness of thin vertical wall can be good index of machining precision. A design depicted by Fig. 2a was used as that for the test specimen employed in this research and the thickness at the center of the wall, which is denominated by “t,” wasmeasured. Laser is scanned in the longitudinal direction of the wall, and we define thedistance between the central lines of the outmost scans as the nominal thickness of the wall, tn (Fig. 2b). Building condition as displayed in Tbl.

2b). Building condition as displayed in Tbl. II is employed for the specimen preparation.Relationships between measured thickness and nominal thickness when various spotdiameters were used are shown in Fig. 3. The spot diameter of 300µm was obtained bydefocusing the optics of scanner B. Though reducing spot diameter from 300µm to130µm or 150µm improved fineness quite clearly, difference between 130µm and150µm is negligible.Though all build parameters except nominal thickness and spot size were the same inthe previous experiments, they should be modified according to the spot size. In thefollowing experiments, each parameters except layer thickness is adjusted so that its nominalization by spot size is equalized. With respect to the spot sizes of 150µm and 300µm, wall thickness measurements were performed using build parameters as shown in Tbl. III.

With respect to the spot sizes of 150µm and 300µm, wall thickness measurements were performed using build parameters as shown in Tbl. III. The same layer thickness was used for the both spot diameters because reducing the layer thickness from the current value of 100µm is difficult or impossible due to difficulty of coating thin layer. As depicted in Fig. 4, relationships between actualand nominal wall thicknesses in normalized form for two diameters are in very goodaccordance.Tbl. III Build parameters of which normalization by spot diameter areequalizedSpot diameter 150μm 2.0m/s Scan speed 50μm Scan spacing 100μm Layer Laser power 1.2W Part bed temp. t m−4K300μm 4.0m/s 100μm 100μm 4.8W t m−4K4.00s s e n k c i h t3.00Spot diameters150ml l a w300ml a u t c a d e z i l2.001.00m r o N00123Normlized nominal wall thicknessFig.

4 Relationships between normalized measured thickness andnormalized nominal thicknessPARAMETER OPTIMIZATION FOR THICKNESS MINIMIZATION By optimizing the process parameters to minimize the wall thickness (Tbl. IV), a wallas thin as 180µm was obtained. The wall was so fragile that even removing unsinteredpowder around the part can easily destroy it. In other words, the minimum wallthickness is dependent on the level of skills in powder removal (breakout process) aswell as build parameters such as laser power. However, standardizing the breakoutprocess to remove unsintered powder necessarily and sufficiently is very difficult.Moreover, building such a weak part does not make sense. We will discuss the strengthof the thin wall in the following section.Tbl. IV Build parameter for the thin wall with the minimal thicknessNominal thickness 165µm Scanning velocity 3.8m/s 100µm Beam diameter t m−4K Laser power130µm Layer thickness 0.43W Powder bed tempFig.

5 Vertical wall of 180µm in thicknessDENSITY OF THIN WALL The mechanical strength of a part that is fabricated by an AM techinique is dependenton its build parameters while the strength of a part obtained by subtractive processesmaintains; subtractive processes preserve the strength, and AM process decides that.Thus, we should discuss relationship between strength of miniature structure (thin wallin this case) and build parameters, but it is challenging or almost impossible to measure the strength of such a small structures. Since strength and density correlate very closely as shown in Fig. 6, we can use the packing rate as an index for the strength and refer this relationship as the need arises. To examine the packing rate of miniature parts specimens as depicted in Fig. 7 were prepared using build parameters as shown in Tbl. V. Fig. 8 shows relationship between packing rate and nominal width of specimens.

7 were prepared using build parameters as shown in Tbl. V. Fig. 8 shows relationship between packing rate and nominal width of specimens. Packing rate was calculated by the following equation,(cid:2010) (cid:3404)(cid:1865)(cid:2929) (cid:1874)(cid:2929) (cid:2025)(cid:2930)(1)60] a P M50[ h t g n e r t s4030e l i s n e T20100.700.750.800.85 Packing rate0.900.951.00Fig. 6 Relationship between tensile strength and packing ratetnm m 0 1Build Direction10mm Fig. 7 Test specimen for density measurementTbl. V Build parameters of specimen for density measurementLaser power Scaning velocity Scaning inberval Layer thicknessPowder bed temperature1.4W 2.0m/s 40µm 100µm t m-3K1.00.9e t a r0.8g n i k c a P0.70.60.50.4012 Nominal tickness [mm]34Fig. 8 Relationship between tensile strength and packing densityHere, ms and vb can be obtained by measuring the specimen.

8 Relationship between tensile strength and packing densityHere, ms and vb can be obtained by measuring the specimen. True density of the powder material, ρt,=1.03g/cm3, which is provided by the powder fabricator. As displayed in Fig. 8, packing rate increases as nominal width is increased. The rate of increase is steepwhen the width is smaller than 1mm. No specimen with a nominal thickness thinnerthan 0.16mm was strong enough to stand breakout process.Packing rate is also affected by supplied laser energy. Fig. 9 shows relationshipsbetween packing rate and laser energy supplied to a unit area when nominal widths wereset to 0.16mm, 0.48mm and 0.96mm, respectively. The packing rete increased as thesupplied energy becomes larger, and it was saturated with the energy is greater than0.900.80e t a R g n i k c a P0.700.60Nominal thickness0.500.16mm 0.48mm 0.96mm0.4010305070Laser energy per unit area [kJ/m 2] Fig.

9 Relationship between packing rate and supplied laser energy forvarious nominal thicknesses2.50Nominal thickness]m m2.000.16mm 0.48mm 0.96mm[s s e n k c i h t1.501.00l l aW0.500 10305070Supplied energy per unit area [kJ/m 2] Fig. 10 Relationship between wall thickness and supplied laser energy forvarious nominal thicknessesaround 25kJ/m2, when nominal width was thicker than 0.48mm. Contrary, saturation was not observed when nominal thickness is 0.16mm. The actual wall thickness alsoincreased as energy supply is increased (Fig. 10).Scan interval is also affecting packing rate. Though the scanning interval of 80µmprovided the largest packing rate(Fig. 11), the difference from narrower intervals werevery small. The difference between 40µm and 20µm is negeligible (Fig. 12).The parameters that can maximize the packing rate with the wall thickness below 1mmwas searched for, and a set of parameters as shown in Tbl. VI was obtained.

12).The parameters that can maximize the packing rate with the wall thickness below 1mmwas searched for, and a set of parameters as shown in Tbl. VI was obtained. This set ofparameters gave 82% of packing rate with a wall thickness of 0.95mm.Fig. 13a is a cross sectional view of this wall, which is obtained by micro CT (SMX-130CT-SV, SHIMADZU) observation. We can find that pores are distributed unevenly, and porosities in the middle and outer portions (Fig. 13b) are 11% and 25%, respectively. Since packing rate was saturated around 80% in overall tendency, parameters that could minimize the wall thickness with packing rate over 80% was searched, and a set of parameters as shown in was obtained. By referring the relationship between strength and density as shown in Fig.

13 Micro CT graphsTbl. VII Build parameters for thinnest wall denser than 80%Nominal thickness Laser power per unit area Scaning inberval Actuatl thickness Packing density0.32mm 31kJ 60µm 0.62mm81%EXPANSION OF BUILD ENVELOPE To expand scanning range, a commercially available laser marker (scanner A) was seton X-Y scanner that was driven by a stepper motor and ball-screw system. To examinethe sintering quality at the seam, a specimen as displayed by Fig. 14 was prepared.Since the region that was sintered earlier is cooled while the scanner is conveyed to thenext region, insufficient sintering occurs at the seam. As a result, porosity at the seambecame 17% while it was 12% at the other place. To compensate this loss of the heatduring X-Y scanner operation, exposure for the later region is overlapped on thepreviously exposed area. Overlapping one scan line could reduce the porosity to 14%,but this is still larger than at non seam cross section.

Overlapping one scan line could reduce the porosity to 14%,but this is still larger than at non seam cross section. Contrary, overlapping two scanlines results in yellowing. Though there must be optimal parameter of overlapping thatcan avoid both of oversintering and undersintering, it is not unique since the loss of heatis dependent on various conditions such as geometry of built parts, place of part inbuilding chamber.In the previous cases, the seaming was performed at the same place in all the layers.This seems to emphasize the effect of undersintering. To disperse the effect, seamingpositions were varied by layers. More concretely, two lines are selected as candidates ofseams, and seaming position were varied by switching them alternately as depicted byFig. 15.

More concretely, two lines are selected as candidates ofseams, and seaming position were varied by switching them alternately as depicted byFig. 15. As a result, the undersintering disappeared and porosity was successfullydecreased to 12%.DISCUSSIONS In this research, investigation on fabrication of precise structure with plastic lasersintering is carried out on an experimental basis. Effect of reducing the spot size of laserbeam is examined by varying the spot size. Two laser sintering apparatuses were10mm25mm2mm25mmFig. 14 Specimen for seam quality examinationActive boundaryBoundary candidatesBuild directionFig. 15 Dispersion of undersintering effect at the seamprototyped to use smaller laser spot than commercially available machine can provide.Regenfuss’s group took similar way to improve the resolution of laser sintering fabrication of metal and ceramic [3].

They used a very small spot diameter of 14µm [4] and obtained very thin vertical wall of 40µm in thickness with a high aspect ratio of 25. Their device was commercialized already [5]. There are many differences betweenmetal and plastic. First of all, we cannot apply short wavelength laser to plastic lasersintering because of its low absorption to regular plastic while most metal laser sintering or melting devices are using YAG or fiber lasers with short wavelengths around 1µm. Since diffraction limit is proportional to wavelength, focusing laser spot for plastic lasersintering into small spot forces us a sacrifice of scanning range as mentioned in theearly section of this paper. Secondly, viscosity of molten plastic is very high and variesvery gently with temperature, while a metal melts at its melting point very quickly andviscosity reduces dramatically.Though this research is carried out on experimental basis, theoretical approach is alsovery important.

With respect to metal sintering or melting, various theoretical modelswere introduced [6...8] in the early stage of its development. However, behavior ofplastic is very much complicated, and recent experiment base report about plasticbehavior or characteristic still can provide new and important information [10].This paper concludes that minimum resolution of plastic laser sintering is around0.6mm. It is obvious this resolution is not acceptable as the performance of an RM tool.We have some room for miniaturization of spot size, but large grain size or layerthickness, we guess, is prevent further improvement in resolution of plastic lasersintering. Thus our important challenge is progress in powder handling. In addition,decrease of focal length which is forced by spot size reduction leads to lowering theceiling of process chamber.

Thus our important challenge is progress in powder handling. In addition,decrease of focal length which is forced by spot size reduction leads to lowering theceiling of process chamber. This is also troublesome from a practical viewpoint.CONCLUSIONS With respect to minimum thickness of walls, reducing spot diameter to 150µm is quit effective as far as we tested in this research. Since there was no significant difference between 130µm and 150µm, we can guess that precision improvement only by reducing spot size comes to the end. A packing rate of thin part is saturated with about 80%, and this gives tensile strength of 30MPa. Though this strength is 60% of that for bulky sinter, it is enough strong for some applications. As the minimum wall thickness with this rate, 0.6mm was obtained.

provement of Electrical Conductivity of SLS PEM Fuel Cell Bipolar Plates Ssuwei Chen, David L. Bourell, Kristin L. Wood Department of Mechanical Engineering, Laboratory for Freeform Fabrication The University of Texas at Austin, Austin, TX 78712Abstract Reviewed, accepted August 3, 2005 Previous work in this research demonstrated the feasibility of fabrication of proton exchange membrane (PEM) fuel cell bipolar plates by an indirect selective laser sintering (SLS) route. Properties of the SLS bipolar plate, such as flexural strength, corrosion resistance and gas impermeability, etc. are quite promising and satisfactory. However, initial results showed that there was still room for the improvement in electrical conductivity.

are quite promising and satisfactory. However, initial results showed that there was still room for the improvement in electrical conductivity. This paper summaries the strategies investigated in an effort to increase the electrical conductivity, among which are: (1) infiltration of brown parts with conductive polymer (2) addition of a liquid phenolic infiltration/re-curing step prior to final sealing and (3) reduction of glassy carbon resistivity by curing process parameter control. Results show that the electrical conductivity value may be improved from 80 S/cm to around 108 S/cm, which is equivalent to a 35% jump, when the phenolic infiltration/re-curing step is applied before final epoxy sealing.Background and Introduction Fabrication of PEM fuel cell bipolar plates by indirect SLS has shown its feasibility based on previous experience [1]. To further increase the plate’s electrical conductivity, three potential remedies were investigated.

To further increase the plate’s electrical conductivity, three potential remedies were investigated. Since most of the strategies explored involved in using infiltration techniques for electrical conductivity improvement, their background and principle theories, Washburn and Young’s equations, will be presented first.The Washburn equation, which is commonly used to obtain pore size distributions by the mercury porosimetry approach, provides the required pressure P to infiltrate a pore of certain radius r . It is expressed as γ= 2(θ γ is the surface tension of liquid-vapor interface, θ is the contact angle between where LV liquid and solid interface and r is the capillary radius. Equation (1) can further be modified to give a relationship between height of rising liquid front h and the corresponding time required t.

Equation (1) can further be modified to give a relationship between height of rising liquid front h and the corresponding time required t . [2].Pcos/)r(1)LVh2 =γ r cos L η 2θ t(2)where η is the viscosity of the infiltrant.In the case of pressureless and spontaneous infiltration of liquid phases into porous preforms, i.e. there is no aid expected from externally applied pressure or vacuum, one of the most critical issues is the adhesion or wetting characteristics of the infiltrant and458substrate. According to Young’s equation, the surface energies at the three phase contacts are related to the equilibrium contact angle through:θ (3) γ are the surface free energies between solid and vapor, solid and where liquid and liquid and vapor phase, respectively.

For spontaneous wetting to occur, therefore, one of the rules relevant to effective bonding low surface energy substrates is: <γ SV=+ γγcosSLLVγ ,γ andSVSLLVγ inγ(4)substratefiltrantTable 1 [3] lists surface energy of some of the common adhesive/infiltrant and substrate materials at room temperature. For obtaining best wetting behaviors and strong γ adhesion forces, the condition: is always desirable. It is evident from in Table 1 that the surface energy of liquid epoxy resin is less than that of graphite. Thus spontaneous infiltration of liquid epoxy resin into a porous SLS brown part is expected to occur. Lastly, the time required for a complete infiltration of porous SLS bipolar plate with epoxy resin can be estimated from rearranging equation (2):γ>>substratefiltrantt=η 2 h γ cos r L2θ=2×(10g.

This necessary step was taken simply to make the porous brown part gas impermeable. It showed that with an adequate viscosity, epoxy resin was able to penetrate into the pore channels inside the part to make it gas tight upon curing. However, our attention at this point was focused on searching for an infiltrant capable of sealing the porous part as well as improving the electrical conductivity simultaneously. Based on the extensive literature survey, it is realized that epoxy resin containing carbon black makes up the greater459portion of conductive epoxy resins available since this additive is the most cost effective; even though some other fillers such as copper, silver and carbon fibril, etc. are available.Theory Electrically conductive compounds made of an insulating polymer matrix and conducting particles have been extensively used in antistatic materials, electromagnetic shielding and industries.

The electrical conductivity of such composites depends significantly on the concentration of the conducting phase and the extent of its continuity. Various mechanisms have been proposed to describe the electrically conducting behavior in polymeric composite materials [4, 5], among which percolation theory and the quantum mechanical tunneling effect are two popular models to illustrate the electron transport process [6]. The law of percolation theory states that a critical concentration or volume fraction, referred to as percolation threshold, of conductive filler is necessary to initiate the electrical conductivity in polymeric composite materials. The material behaves as an insulator when the filler concentration is too low to form a connecting network of conductive sites.

The material behaves as an insulator when the filler concentration is too low to form a connecting network of conductive sites. On the other hand, the material undergoes a sharp transition from nonconductor to conductor when the filler concentration is above the percolation threshold, normally around 7% [7], allowing electrons to tunnel through or jump between closely distributed filler sites. Figure 1 [6] demonstrates an example of the experimental results of the relationship between electrical resistivity and carbon black (CB) fraction for CB filled high-density polyethylene (HDPE) compound.liquid crystal display (LCD) assemblyFigure 1. Electrical resistivity and carbon black fraction relationship for HDPE/CB compound. (adapted from Tang et al.

Electrical resistivity and carbon black fraction relationship for HDPE/CB compound. (adapted from Tang et al. [6])Experimentation and Result To assess the feasibility of electrical conductivity improvement on SLS bipolar plate by using conducting epoxy resin infiltration, we would like to first know the maximum achievable electrical conductivity value a cured conducting epoxy resin can460have. A commercially available epoxy resin and hardener obtained from System Three Resins, Inc. were mixed with carbon black powder purchased from Alfa Aesar Company. The epoxy was Clear Coat Epoxy Resin mixed with the hardener in a ratio of 2:1. The carbon black powder has an average particle size of 0.042 micron and a specific surface area of 75 m2/g. Ten cups of carbon black loaded epoxy resins were prepared and cured overnight in an oven to form ten solid samples. The carbon black loading level for each sample was 1, 2, 3, 5, 7, 10, 12, 15, 20 and 25 volume percent, respectively.

The carbon black loading level for each sample was 1, 2, 3, 5, 7, 10, 12, 15, 20 and 25 volume percent, respectively. Each cured solid sample was then cut into four 3 mm by 3mm by 30 mm test specimens, making a total of 40 specimens, for electrical resistivity testing according to ASTM D257 specifications. This four point probe technique measures electrical resistivity by applying a constant current over the test specimen and measuring the voltage drop across the specified distance. The electrical resistivity can subsequently be calculated by following Equation (5):wtV∆ ( )( ) )( LI where γ is electrical resistivity, V∆ is voltage drop across a prescribed distance, t is specimen thickness, w is specimen width, L is the distance over which V∆ is measured and I is applied constant current. A Keithley 224 Programmable Current Source and Keithley 181 Nanovoltmeter were used to perform the tests.

A Keithley 224 Programmable Current Source and Keithley 181 Nanovoltmeter were used to perform the tests. Figure 2 shows the average resistivity value for cured conducting epoxy resin versus its carbon black loading.=γ(5)Figure 2. Resistivity of CB-filled epoxy resin versus CB content. The first two circle points are predicted values.As can be noticed from the plot, the electrical resistivity curve is expected to drop dramatically from nonconductive to conductive behavior over a narrow range of carbon black loading due to the development of a network of closely-seated CB particles. The predicted very high electrical resistivity for 1 and 2 vol % CB specimens was not obtained due to the limits of our equipment. More importantly, the minimum resistivity461achieved is around 105 Ω-cm, which is equivalent to a maximum electrical conductivity value of 10-5 S/cm when CB loading is roughly above 8 vol. %. The result is very similar to what is reported elsewhere [8, 9].

%. The result is very similar to what is reported elsewhere [8, 9]. However, the outcome is not promising for our application since it is far below our current conductivity value and does not seem to help in the overall conductivity improvement to a value of over 100 S/cm. What was even more disappointing is that the fine carbon black powder tends to agglomerate easily in the CB-filled epoxy resins and largely increase the resin viscosity. High resin viscosity and CB agglomeration can block CB powders from flowing into pores of the porous plate and limit the performance of infiltration. Figure 3 presents the photos showing that the agglomeration and sedimentation of carbon black powder in epoxy resin become more severe as CB loading increases. It is apparent, at this point, that infiltration of conducting epoxy resin will not be workable for the conductivity improvement of our SLS plates.

It is apparent, at this point, that infiltration of conducting epoxy resin will not be workable for the conductivity improvement of our SLS plates. Even many commercially available conducting epoxy resins exhibit much higher values of conductivity; most of them are in the form of “paste” instead of what is required here, a low viscosity infiltrant.Figure 3. Agglomeration and sedimentation of CB powders in epoxy resins. (Left) 1 vol. % of CB loading. (Right) 15 vol. % of CB loading.Liquid Phenolic Infiltration/Re-Curing The electrical conductivity of powder metallurgy parts is affected by many variables, such as impurity, pore shape, and connectivity, etc, and one of the most significant factors is the amount of pores presented in the solids. Many models for thermal and electrical conductivity, based on theoretical or semi-empirical results, have been reported [10-13].

Many models for thermal and electrical conductivity, based on theoretical or semi-empirical results, have been reported [10-13]. From these models, it is obvious that electrical conductivity increases as the percentage of porosity is reduced while the part becomes denser. This trend brought up the idea for improving electrical conductivity of SLS bipolar plate by increasing the glassy carbon residue level so as to replace portion of the porosity space that is originally occupied by the non-conducting epoxy resin when final epoxy sealing is completed.462Phenolic resin was previously chosen as one of the constituents along with graphite particles to make up the powder mixture for the requirements of this application. It is selected here as the intermediate infiltrant to elevate the carbon residue level when recuring process is executed. The phenolic resin GP5546 PARAC Powdered Phenolic Resin was obtained from Georgia Pacific Resins, Inc.

The phenolic resin GP5546 PARAC Powdered Phenolic Resin was obtained from Georgia Pacific Resins, Inc. and was dissolved in acetone at a temperature of 50 ºC by using a magnetic stirrer for 30 minutes. The phenolic infiltrant was made of 60 vol. % of powdered phenolic resin and 40 vol. % of acetone for the first infiltration and the phenolic infiltrated brown part was placed back in the high temperature furnace for recuring. The recuring process was performed under inert gas atmosphere as the furnace chamber was filled with research grade argon gas from PraxAir Inc. The temperature profile for the recuring cycle started at room temperature and went up to 800 ºC with a 1 ºC/min heating ramp rate and was held for 1 hour before cooling back to room temperature naturally. The recured brown part was less porous compared to brown parts that did not receive such treatment. Electrical conductivity was improved to an average value of 108 S/cm.

The recured brown part was less porous compared to brown parts that did not receive such treatment. Electrical conductivity was improved to an average value of 108 S/cm. Following the same strategy, a second phenolic infiltration/recuring step was performed in an attempt to further increase the conductivity. The phenolic infiltrant was modified to a composition of 45 vol. % of powdered phenolic resin and 55 vol. % of acetone for reduced viscosity. The same temperature profile was implemented for the second time recuring process and the final electrical conductivity was increased to an averaging value of 117 S/cm. Figure 4a and Figure 4b are SEM micrographs of a brown part with and without a phenolic infiltration/recuring step, respectively. It can be seen that the amount of carbonized phenolic resin had increased by way of the addition of this fabrication step compared to that without phenolic infiltration/recuring treatment.Figure 4(a).

SEM micrograph of part with a phenolic infiltration/recuring step. The dendrite-like particles are glassy carbon while the larger flakes are graphite particles.463Figure 4(b) SEM micrograph of part without a phenolic infiltration/recuring step. Less glassy carbon residuals are able to be observed.Figure 5 is a modified flow chart for SLS bipolar plate fabrication with added phenolic infiltration/recuring steps.Figure 5. Modified flow chart for bipolar plate manufacturing using SLS.In Figure 6, a bar chart for electrical conductivity with respect to each fabrication step is demonstrated. As can be seen, there is a 35% boost in electrical conductivity when the first phenolic infiltration/recuring step is employed and another 8.3% increase when the second phenolic infiltration/recuring step is done.

The reduced rate in electrical conductivity improvement implies that less glassy carbon residues are deposited each time inside the brown part since the corresponding amount of porosity is lessened.464Figure 6. Electrical conductivity improvement of SLS bipolar plate with respect to corresponding fabrication step.Curing Process Parameter Control The brown part is composed of graphite particles and glassy carbon produced from carbonization of phenolic resin, and both ingredients contribute to its final conductivity. If the conductivity of glassy carbon were able to be significantly improved, it might be able to benefit the overall conductivity of brown part supposing that the conductivity of graphite particles is constant. Bhatia et al. [14] explored the variation of some physical properties of carbonized phenolic resin when increasing the curing temperature, one of which gives the relationship between its electrical resistivity and curing temperature as shown in Figure 7.

It is evident from the curve that the electrical conductivity of pyrolyzed phenolic falls drastically in the range of 600º C to 800º C, beyond which the rate of fall in resistivity decreases gradually.To examine the extent of the effect of conductivity increase from glassy carbon, the peak dwell temperature for brown part formation was raised from the original 800º C to 1000º C and electrical conductivity values were measured by using the four point probe technique described previously. Results showed, unfortunately, that the electrical conductivity of brown part does not exhibit significant improvement from the increase in final curing temperature. This may be attributed to the fact that only a relatively small portion of glassy carbon exists in the composition and the most critical factor dominating the electrical conductivity in powder sintered parts is its porosity level.

Electrical conductivity of powder sintered parts normally follows a power law and is a strong function of porosity level inside the part [13, 15].465Figure 7. Variation of electrical resistivity of cured phenolic resin versus temperature. (from Bhatia et al. [14])Conclusions Three potential strategies were examined to ascertain the most appropriate method for improving the electrical conductivity of SLS bipolar plates. Inclusion of liquid phenolic infiltration/recuring steps to the fabrication route proved to be the most effective solution among others. This method is very similar to the way people used to make carbon/carbon or carbon/graphite composite materials where multiple cycles of thermosetting resin impregnation and curing are performed to form a gas-tight matrix.

For our special application of fabrication of SLS PEM fuel cell bipolar plates, however, simply one or two times of liquid phenolic infiltration/recuring step are necessary to meet the electrical conductivity target value, especially when considering the lengthy and costly polymer pyrolysis process. The treated part is subsequently infiltrated with liquid epoxy resin to seal its surfaces. One other possible way to improve the electrical conductivity is chemical vapor infiltration (CVI) of carbon from hydrocarbon gaseous precursors into porous SLS brown parts. This technique is potentially capable of creating a hermetic skin on the plate surfaces as well as improving electrical conductivity due to the addition of elemental carbon.466References [1] S. Chen, D. L. Bourell and K. L. Wood “Fabrication of PEM Fuel Cell Bipolar Plate by Indirect SLS” The 15th Solid Freeform Fabrication Symposium Proceedings, August (2004), pp. 244-256 [2] A. Siebols, et al.

[15] J. Koh and A. Fortini, “Prediction of Thermal Conductivity and Electrical Resistivity of Porous Metallic Materials,” International Journal of Heat and Mass Transfer, Vol. 16 (1973), pp. 2013-2022.Acknowledgement This research was funded by State of Texas Technology Development and Transfer Grant Number 003658 and by the US Office of Naval Research Grant Number N00014-00-1- 0334.467.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperImprovement of recycle rate in laser sintering by low temperature processT. KIGURE*, T. NIINO†Tokyo Metropolitan Industrial Technology Research Institute †Institute of Industrial Science, the University of TokyoAbstractLaser sintering process selectively solidifies its powder bed to obtain designated parts and leaves the rest unsolidified. When the remained powder is recycled, a certain amount of fresh powder is added to moderate the effect of deterioration by preheating in the previous batch. In terms of economy, improvement of the recycle rate of used powder is one of the big challenges. The authors are developing a novel laser sintering process that prevents part being processed from warping not by preheating the powder bed but by anchoring the parts to a rigid base plate.

The authors are developing a novel laser sintering process that prevents part being processed from warping not by preheating the powder bed but by anchoring the parts to a rigid base plate. Since the new process, namely low temperature process, can lower the bed temperature than normal high temperature process, it is expected to reduce deterioration of the used powder. In present research, processablity of recycled powder is evaluated by its MFR, and it is shown that operation at a high recycle rate more than 90% is possible.IntroductionIn laser sintering process, a part of powder bed is selectively solidified by laser irradiation to obtain designated parts. Rest of the bed, the powder that remains unsolidified when a process for one batch is finished, is reused in the following batch.

Rest of the bed, the powder that remains unsolidified when a process for one batch is finished, is reused in the following batch. Exposed to high temperature, the unsolidified powder is deteriorated [1], and using the powder directly in the next batch leads to defects such as rough surface known as “orange peel.” To moderate the effect of material deterioration, used powder is mixed with a certain amount of fresh powder. Since material cost for one batch is equivalent to the cost for added fresh powder, the amount of the additional powder directly impacts material cost of part production. Thus, reduction of additional material is a big challenge in practical use of laser sintering technology. Generally, rate of additional fresh powder, refresh rate, in feed stock is 30% to 50%. [2, 3] On the other hand, in most cases, the volume occupied by parts in a bed is only 10% or less of the whole powder bed due to difficulty in part arrangement.

[2, 3] On the other hand, in most cases, the volume occupied by parts in a bed is only 10% or less of the whole powder bed due to difficulty in part arrangement. Thus, amount of the powder to be purchased becomes three to five times of part volume. If repeated use of material at refresh rate of 10% can be, it becomes the amount of newly added material is almost equivalent to the amount of material consumed as a product. In deterioration of powder during process, powder bed preheating plays a dominating role. In laser sintering, powder bed temperature is maintained in a range between melting and recrystallization temperature, which is called “process window.” Within the process window, the melted material stays in supercooled liquid state so that thermal stress that can be a cause of part warpage is suppressed. On the other hand, this powder bed preheating causes deterioration of the material.

On the other hand, this powder bed preheating causes deterioration of the material. The authors are introducing a novel laser sintering process that prevents parts from warping by anchoring them to a rigid base plate. Since the process, namely “low temperature process,” does not utilize supercooling phenomenon, it is not necessary to maintain the powder bed temperature within the process window during process. This unnecessity of tight temperature control [4-6] allows process temperature to be less than crystalizing temperature, and low bed temperature prevent the material deterioration leading to improvement of recycle rate. The aim of this research is to access effectiveness of low temperature process in improvement of powder recyclability. Recycle powder with a refresh rate of 10% was tested. Several times of build tests were performed repetitively as the powder is regenerated after each build. Performance variation of powder and build parts are evaluated.

Several times of build tests were performed repetitively as the powder is regenerated after each build. Performance variation of powder and build parts are evaluated. MFR of powder material was also measured after each build to evaluate degradation of powder. For evaluation of part quality, relative density and surface roughness were used.550Material and MethodMaterial PA12 powder (Vestosint®, daicel-evonik) was employed. This powder is developed for use in laser sintering. Serial tests started with brand-new fresh powder. After each test is finished, used powder from the bed which was collected from only part-cake was sieved and mixed with brand-new powder at a mixture rate of 9:1 to obtain powder stock for the next batch. DSC measurement result of the powder material is shown in Figure1. The melting point is 184 °C, the peak temperature of recrystallization is 144 °C.

DSC measurement result of the powder material is shown in Figure1. The melting point is 184 °C, the peak temperature of recrystallization is 144 °C. True density of the material is 1.03g/cm315-10-5051015050100150200250300Heat flow [mW]Temperature [oC]Figure 1. DSC curve of vestosint® fresh powder (Heating rate: 10°C /min. Holding at 260°C for 5 min. Cooling rate: 5°C /min .)Process Parameters A commercially available machine (RaFaEl300C®, ASPECT) was used. Process parameter set is shown in Table 1. The parameters of low temperature process are obtained by the preliminary test as those that provide the highest density as a primary index of mechanical strength. Powder bed temperature of 130°C for low temperature process is lower than the recrystallization temperature of the powder. Table 2 shows building time and data height of each process in this research. The building time was set to be roughly the same between low temperature process and high temperature process.

Table 2 shows building time and data height of each process in this research. The building time was set to be roughly the same between low temperature process and high temperature process. Each process was repeated 5times. Figure 2 shows build layout of each process. In all builds, parts were arranged so as to minimize mutual thermal effect.Building processTable 1. Processing parameter set Laser powerPowder bed temperatureLayer thicknessScan intervalScan speedEnergy per unit areaLow temperature High temperature130oC170oC0.1mm0.1mm0.10 mm0.14 mm 55112.0W24.5W2.0 m/s60.0 mJ/mm210.0 m/s 17.5 mJ/mm2Table 2. Build time and data height of each processNumber of cycles1 2 3 4 5Low temperature process High temperature processTime 3h06min 3h05min 3h22min 3h00min 3h00minHeight 9mm 9mm 9mm 9mm 9mmTime 3h13min 2h28min 3h11min 2h39min 2h02minHeight 40mm 28mm 40mm 28mm 18mma)Base plateb)Heat barrierFigure 2.

Build layout a) low temperature process b) high temperature processBase plate A base plate to anchor the parts on was employed to prevent them from warping. It was a flat plate of the same material manufactured by high temperature process. To ensure the flatness of the anchoring surface, it was back-upped with rigid aluminum palate which was tightly fixed to the plate by screws as shown in Figure 3.PartsBoltSupport structureBase plateBack up plateFigure 3. Base plateParts density Since relative density of a parts has a strong correlation with its mechanical strength, part quality provided by each powder condition was evaluated by the density.

Base plateParts density Since relative density of a parts has a strong correlation with its mechanical strength, part quality provided by each powder condition was evaluated by the density. Densities of cubic specimens with nominal dimensions of 80mm×10mm×4mm were measured by Archimedes method using following equation. parts =m airm air - m water  water,where mair is weight of a part in air and mwater is weight in water, and ρwater is density of the water.552Surface roughness Since rough surface of parts is a major defect often found when we use degraded material, surface roughness of upper surface of parts was measured. Using 3D laser microscope (OLS4000, Olympus), surface profile in the range of 1.8 mm × 1.8 mm was measured at a magnification of 432 times. Arithmetic mean height Sa was used as evaluation index of surface roughness.Melt Flow Rate (MFR) It is well known that MFR of powder material decreases as quality of the material is degrading [1, 2, 7].

Unsolidified powder was collected from the part-cake after process, and MFR was measured by melt flow tester (MELT INDEXER P-101, Toyo seiki) at a load of 2.16kg and temperature at 235°C.Experimental resultDensity Figure 4 shows the relative densities for each process cycles. Density was higher in low temperature process for all the process cycles. In the both of low temperature and high temperature processes, density decreased as process was repeated in general. The decrease rate of the relative density obtained by low temperature process was smaller.Low temparetureHigh temparetureRelative density [%]Build cycle(130oC)(170oC)909294969810012345Figure 4. Relative density of each build cyclesSurface roughness Figure 5 shows the measured arithmetic mean height Sa. Roughness of part from high temperature process showed larger irregularity and saturates after cycle 3. Roughness of low temperature processed part increased until cycle 4 and decreased at cycle 5.

Roughness of part from high temperature process showed larger irregularity and saturates after cycle 3. Roughness of low temperature processed part increased until cycle 4 and decreased at cycle 5. For the process roughness was doubled after several processes.553051015202530354012345High temparetureArithmetical mean height of the surface Sa [m]Build cycle(130oC)(170oC)Low temparetureFigure 5. Arithmetic mean height Sa of each build cyclesMelt Flow Rate Figure 6 shows variation of MFR. Low temperature reduced degradation of MFR in general. MFR degraded very much after 1st cycle, but decrease became negligible after cycle 3. MFR of powder material used in low temperature process decreased by a factor of 35% while MFR decrease of high temperature processed powder was 52%.Low tempareture(fresh powder)High temparetureMFR [g/10min]Build cycle(130oC)(170oC)0510152025303540012345Figure 6.

MFR of unmetled powder in each build cycles 554DiscussionReduction of part density as an index of mechanical property was less than 1% after processed 5 times. This result shows that recycle rate of 10% enough for low temperature process in terms of mechanical performance. Contrarily, density drop of high temperature process was greater than 2%. Surface roughness was doubled after processed 5 times for the both of low and high temperature processes. Effectiveness of using low temperature process in variation of surface roughness is not clear. However, absolute value of surface roughness obtained by low temperature process was a much smaller than high temperature process for all process cycles. This reason is assumed as follows. Surface of a melted part of bed is relatively smooth owing to surface tension and gravity force.

This reason is assumed as follows. Surface of a melted part of bed is relatively smooth owing to surface tension and gravity force. However, in high temperature process, the surface is roughen again when additional powder is poured for the next layer since it is adhered on the surface which is staying unsolidified in supercooled state. On the other hand, in the low temperature process, the surface once flattened by the tension and the gravity force does not adhere additional powder since the surface solidifies as soon as laser irradiation is finished. The increase in surface roughness due to process repeating may be due to the increased viscosity of the melt as Wudy et al. suggested [8]. The decrease in MFR became negligible after cycle 3. This is similar trend to 32% and 50% refresh rate in the high temperature process and indicate effect of mixing refresh material [3,7]. In the MFR of high temperature process in this research, there is no effect of mixing of fresh materials.

In the MFR of high temperature process in this research, there is no effect of mixing of fresh materials. Degradation of MFR after process cycles is smaller for low temperature process than high temperature process. This indicated that powder bed temperature lower than recrystallization temperature suppressed effect of heat in material deterioration. According to the heating experiment by Dotchev et al., they are shown that the value of MFR clearly differs depending on the temperature and higher MFR leads to better product quality [2]. This support that low temperature process can reduce material deterioration, and the better part quality derives this suppressed material damage.ConclusionIn this research, improvement of material recycling rate of laser sintering using low temperature process was discussed. The process was repeated five times powder was regenerated at a refresh rate of 10%.

The process was repeated five times powder was regenerated at a refresh rate of 10%. Degradation in part density as an index of material property became less than 1% after the repeated processes and saturated. In terms of mechanical performance, we can conclude low temperature process is quite effective for improvement of powder recyclability. Surface roughness was increased by a factor of two in the both of low and high temperature process. Although we did not find advantage of using low temperature process in prevention of degradation in surface finish, low temperature process always provided better surface quality than high temperature process. Low temperature process could not suppress drop of MFR after the first process using brand-new powder, but the decrease became negligible after the third process. This shows that 10% of refresh rate is enough in terms of material deterioration for low temperature process.

This shows that 10% of refresh rate is enough in terms of material deterioration for low temperature process. We can conclude that low temperature process can eliminate cost multiplication issue due to low recyclability.AcknowledgementsThis research is a part of “Strategic Innovation Promotion Program” supported by Cabinet Office, Government of Japan.Reference[1] Gornet et al, “Characterization of selective laser sintering materials to determine process stability”, Proc. Solid Freeform Fabrication Symposium 2002 (2002) 546-553 [2] Dotchev et al, “Recycling of polyamide 12 based powders in the laser sintering process,” Rapid Prototyping Journal, 15 (2009) 192-203555[3] Josupeit et al. “Material properties of laser sintered polyamide 12 as function of build cycles using low refresh rates”, Proc. Solid Freeform Fabrication Symposium 2015, (2015) 540-548 [4] Niino et al, “Low temperature selective laser melting of high temperature plastic powder,” Proc.

PROVEMENTS IN SLS PART ACCURACYChristian Nelson Kevin McAlea Damien Gray DTM Corporation Austin, Texas 78754ABSTRACTSLS® part accuracy is influenced by a number of machine and material characteristics. Some of the most significant sources of error are associated with laser beam positioning (static and dynamic) on the part bed surface and uncertainty in the calibration factors used to compensate for material shrinkage and growth as well as the finite width of the laser beam. Another source of error is the minimum resolution of the process, which is a dependent on the particle size and shape of the material. In this presentation, technical background on these issues will be provided. In addition, part data obtained with a number of SLS materials demonstrating improved accuracy obtained through machine modifications and improved calibration methods will be described.1.INTRODUCTIONRapid prototyping part accuracy has been reported in a number of studies.

In the case of stere0 lithography, a large amount of accuracy data has been generated for a model geometry to characterize process and material technology improvements over time (1, 2). More recently, a number of studies which compare the accuracy of different rapid prototyping technologies have been documented (3-6). In nearly all of these studies, accuracy has been reported as the deviation of measured dimensions from the desired dimensions. Other components of accuracy are also important to users of these technologies. In particular, trueness of features as characterized by the sharpness of corners and edges and detail resolution is quite important. as is the ease of part In this study, improvements in the accuracy of SLS parts as characterized by fit and assembly. both dimensional tolerances and trueness of features will be discussed. In addition, a discussion of critical material and process issues which affect accuracy will be provided.2.

both dimensional tolerances and trueness of features will be discussed. In addition, a discussion of critical material and process issues which affect accuracy will be provided.2. BACKGROUNDThe SLS Process A schematic of the SLS process as commercialized by DTM is shown in Figure 1. The key components are a CO2 laser, laser optics and scanning mirrors, part build cylinder, feed cartridges, heaters, and a roller mechanism. To fabricate a part, powder from one of the feed cartridges is first spread in a thin, uniform layer over the part bed surface. A typical layer thickness is 0.12 mm. The laser beam is rastered over the part bed surface with the scanning mirrors and the laser energy modulated so that only the area which corresponds to the cross section of the object is fused. The 2D cross sections are defined by software which "slices" the 3D CAD information into layers. A new layer of powder is then spread from the other feed159processed with consolidation n ..... "·'...."H.

A new layer of powder is then spread from the other feed159processed with consolidation n ..... "·'...."H. . . . about 1%), butconsiderable growth. Crystalline or semicrystalline materials exhibit high shrinkage (typically 3 4%) associated with both the material phase change and the high degree of consolidation obtained during processing, but the amount of growth is usually limited by the relatively low laser powers which are required to completely melt the material. While software is used to compensate for the dimensional changes associated with .material shrink and growth, the best accuracy is usually obtained when these compensations are small. For a given material, it is possible to moderate shrink and growth by adding a reinforcement or filler to the system (8, 9). It is also possible to achieve near zero shrinkage with multiphase coated particle morphologies (10, 8).The primary effect of particle size and shape on part accuracy is on the trueness of features.

Smaller particles with a high degree of sphericity produce parts with sharper edges and corners In addition, the layer thickness of parts can be reduced for and improved detail resolution. smaller particle sizes which reduces the "stair step" effect which is common to all layered manufacturing processes.2.3 Accuracy - Process Effects The key process elements which affect part accuracy are the beam delivery and thermal control systems. As previously described, the beam delivery system uses scanning mirrors to position the laser beam on the part bed surface. There are errors associated with the inherent positional accuracy of the scanning system as well as dynamic errors associated with coordinating the laser "on" and "off' with the acceleration and deceleration of the mirrors. This type of dynamic error causes scan vectors to be drawn with a non-uniform energy density. In addition, the finite size of the laser beam causes material outside the part boundaries to be scanned.

This type of dynamic error causes scan vectors to be drawn with a non-uniform energy density. In addition, the finite size of the laser beam causes material outside the part boundaries to be scanned. Part dimensions are adjusted for the finite diameter of the beam through a software compensation, just as they are for growth. From a practical standpoint, the effects of growth and finite beam diameter on part accuracy are indistinguishable since they both cause a constant increase in part dimensions.Adequate thermal control is necessary to prevent accuracy problems associated with part warpage or "curl". Warpage can occur at the part bed surface or "in-build", or as the part is cooled from the build temperature to room temperature or "post-build". In-build curl is caused by thermal stresses and layerwise shrinkage stresses. These stresses can be minimized by maintaining the part bed at a temperature near the melting/softening temperature of the material.

These stresses can be minimized by maintaining the part bed at a temperature near the melting/softening temperature of the material. At this temperature, thermal gradients between the part and surrounding powder are minimized while for viscoelastic materials such as polymers, stress relaxation is activated. Post-build curl is caused by stresses which develop as the result of non-uniform part cooling. This mechanism is particularly important for high temperature materials. The control of post-build curl through cooling rate control has been previously described (7, 11). An additional thermal accuracy issue can result from point-to-point temperature differences in a consolidated part. Temperature variability is caused by non-uniform power densities experienced by different cross-sections as the laser is scanned across a part geometry. This thermal variability can cause non-uniform shrinkage and growth.

This thermal variability can cause non-uniform shrinkage and growth. Temperature feedback control techniques for eliminating these effects have been discussed (12).The current work is focused on three of the accuracy elements described above: scanning dynamics, scale/offset calibration and compensation software, and particle size and shape control. Improvements in these areas are described.1613.IMPROVEMENTS3.1.1 Scanner Dynamics Inherent with any scanning system are errors associated with the positional accuracy and dynamics of the galvanometers. The magnitude of the errors is affected by the mass of the mirrors and manufacturers design. The dynamic error is defined as the error that results from exceeding the bandwidth of the system, where the bandwidth is how fast the system can accurately respond to changes.

The dynamic error is defined as the error that results from exceeding the bandwidth of the system, where the bandwidth is how fast the system can accurately respond to changes. The impact of dynamic errors on part accuracy can be reduced by slowing down the scan speed, minimizing the mass of the galvanometer mirrors, or changing the design of the galvanometers. DTM is currently evaluating a set of experimental galvanometers with different design parameters than the current galvanometers.. Dynamic errors were measured in a machine using a simple test where two series of marks are burnt onto a sheet of Mylar. The first series is drawn at a slow scan speed (0.86 mls), and the second set is drawn at a fast scan speed (2.36 mls). The marks made at the slow scan speed are used as a reference from which the second series of marks is measured. The dynamic error is the difference between the actual and the ideal position of the marks drawn at the fast scan speed.

The dynamic error is the difference between the actual and the ideal position of the marks drawn at the fast scan speed. The dynamic errors measured for the current and experimental galvanometers are ± 0.25 and ± 0.025 millimeters, respectively, as shown in Figure 2. The results of this test show that the galvanometer design plays an important role in dynamic performance.0.40S S I-l' 0 l:: l:J;.l (,) .§C':l t=l :>-. 0 4-f 0 ~ bI.l0.30.20.100.10.2~~0.30.4current galvanometersexperimental galvanometersFigure 2. Dynamic errors of the current and experimental galvanometers. Errors were measured along thex and y axes across a 12 inch area.3.1.2 Scanner Repeatability Scanner repeatability is a measure of how accurately the galvanometers can return to a given location in the build area. The repeatability specification for the current galvanometers is 120 J.lrad.

The repeatability specification for the current galvanometers is 120 J.lrad. In the Sinterstation 2000, where the distance from the galvanometers to the build plane is approximately 616 millimeters, the galvanometers are162repeatable to within 0.076 millimeters (0.003 inches). The experimental galvanometers have a repeatability specification of 10 J,trad, or 0.006 millimeters (0.00024 inches) at the build plane.Scanner repeatability dictates how well the ends of the scan vectors drawn on the build plane are aligned; the alignment of the vectors affects the smoothness of vertical surfaces. The combined effects of the scanner dynamics. and scanner repeatability influence overall SLS part accuracy. There is a significant improvement in part accuracy when the new galvanometers are used to build and accuracy benchmark part in a polymer coated metal material as shown in Figure 3. This material has essentially zero shrinka.ge in the SLS process.0.40.3S S ,....; c:l .....

This material has essentially zero shrinka.ge in the SLS process.0.40.3S S ,....; c:l ..... l=l S 0 Z S 0 ~ -0.1 d ..... 0 ~..... :> Q.) 00.20.100.20.3- - - - - - - - - - - - - --0.4current galvanometersexperimental galvanometersFigure 3. Error distribution for an accuracy part made using polymer coated metal material. The solid box contains 90% of the data, and the dashed box contains 50% of the data. There are 46 distinct x and y measurements in each data set.The initial evaluation of the experimental galvanometers on different SLS machines at different sites began in May 1995. The evaluation included two builds with the current galvanometers, a retrofit of the scanner system, and the same two builds with the experimental galvanometers. The first build contained 8 parts representing a range of part geometries. These parts were built only once with each set of galvanometers to show the first time build accuracy.

The first build contained 8 parts representing a range of part geometries. These parts were built only once with each set of galvanometers to show the first time build accuracy. Also, the scale and beam offset values, calculated especially for the SLS machine and material, were applied to all of the STL files in the build packet.Preliminary results of the evaluation program are presented in Table 1. Parts built with the experimental galvanometers show less random error than parts built with the current galvanometers. The range of error is reduced by approximately 25 to 35% to a value of ± 0.25 mm for nylon and polycarbonate, while the range of error is reduced by approximately 40 to 50% to a value of ± 0.10 mm for metals. The different range of error values for the polymers and163attributed to some of the materials issues discussed earlier. Parts coated made from the polymers experience errors associated with thermal variability, shrinkage as well as beam positioning.

Parts coated made from the polymers experience errors associated with thermal variability, shrinkage as well as beam positioning. Parts made with the polymer coated metal system, however, are only affected by beam positioning errors since the material is processed at ambient temperature and there is essentially no density change associated with the laser processing.canTable 1.Summary of accuracy measurements from the multiple site evaluation of the experimental galvanometers. The part build contained 8 parts representing a range of geometries.

The optimization is On (La) and the Laser Off (LF) parameters (13). The Laser On command focused on the compensates for the initial delay caused by the inertia of the fast axis mirror and rotor. When the scanner controller sends a ramp signal to the galvanometer, telling it to change position, there is a finite delay before the galvanometer starts to move, as is shown in Figure 4a. Because the galvanometer system does not have position feedback to the controller, the controller sends the ramp signal to the galvanometer then waits for a time interval to elapse, specified by the parameter, before turning the laser ON.the vector scan, the scanner still lags the ramp signal by some amount and requires At the end scanner to Off delay allows the actual endpoint of the drawn vector. time to In Figure 4b, the ramp signal is reach the end of the drawn vector before the laser is their final sent to the Off position until approximately 1500 Ils have elapsed.

time to In Figure 4b, the ramp signal is reach the end of the drawn vector before the laser is their final sent to the Off position until approximately 1500 Ils have elapsed. The time interval specified by the parameter starts when the stop command is sent to the galvanometer.time equal zero, but the galvanometers do not164The optimal set of LO and LF parameters are a function of the scan speed. As the scan speed increases, the Laser On and Laser Off delays decrease because the galvanometers accelerate/decelerate more rapidly. To optimize the parameters, high and low values are chosen for the LO and LF, then a designed experiment using a structured set of permutations of these parameters is used to build SLS parts.

To optimize the parameters, high and low values are chosen for the LO and LF, then a designed experiment using a structured set of permutations of these parameters is used to build SLS parts. The SLS parts are then measured, and the optimal set of LOILF parameters are calculated on the basis of part accuracy, repeatability, surface finish, and eccentricity.An appropriate range ofLO and LF values can be found by scanning a position sensitive detector (PSD) and plotting the galvanometer position as a function of time, Figure 4. Data collection is initiated when a move signal is senUo the galvanometers, and thus the LO and LF delays can be directly related to the ordinate ofthe graph.

Data collection is initiated when a move signal is senUo the galvanometers, and thus the LO and LF delays can be directly related to the ordinate ofthe graph. Using the data in Figure 4, at a Step Size of 80, the initial range of values used for the LO and LF were 700 to 1300Jls and 11 00 to 1700Jls, respectively.3.53.544~ c:f 0." ~4.55~ .~ .;;; rf.4.555.55.56~~~_705001000.1500200060---------~-~-~-~~ 500 2000---------~-~-~-~~ 500 time, ItsFigure 4. Galvanometer position as a function of time (left: acceleration curve; right: deceleration curve). The galvanometer is moving at a 8tep Size (88) of 80 LSB's. A move signal is sent to the galvanometers at time equal zero. Legend: 0100 mm vector,. 10 mm vector.3.2 Scale/Offset Calibration and Compensation As discussed in Section 2.2, there are dimensional changes associated with material shrinkage and growth during the fabrication of the SLS object.

Software is available to compensate for these effects, given the proper scale and offset values. DTM has recently revised the procedure to calculate scale and offset values for each of the Laserite materials. The new scale/offset procedure includes the building and measurement of parts, Figure 5, that contain nominal dimensions ranging from 5 to 178 mm (0.2 to 7.0 inches). The deviation from the nominal for each dimension is plotted versus the nominal dimension, and the best linear fit is drawn through the data. The slope of the line is the shrinkage (or expansion) value. A negative slope corresponds to a shrinkage value. The scale value required to compensate for the shrinkage is equal to I/(l-shrink). The intercept of the line is twice the beam offset value (2e!), because all of the dimensions are measured on outward facing surfaces. The beam compensation software moves the skin (or surface) of the STL file inward a specified distance, d.165Figure 5.

The beam compensation software moves the skin (or surface) of the STL file inward a specified distance, d.165Figure 5. Parts built and measured to calculate scale and offset values for the polymer materials used inthe Sinterstation 2000 System.The current methods of beam compensation include normal offset and constant offset. Of these In the normal offset two methods, the normal offset method provides the most uniform results. method, a vector normal is calculated at each vertex based on the normals of the adjacent facets. The vertex is moved along the vertex normal to the new offset position. This method works best for STL files that have circular and flat surfaces because the angle between the facets is large. As the angle between the facets decreases, the distance that the vertex is moved by the normal offset method is not sufficient to move the surface of the STL file by the desired amount. The constant offset method works best for square features.

The constant offset method works best for square features. In this method, the vertex is offset by moving first in the x direction, then in the y direction, by the beam offset value, Figure 6c.A new method of beam compensation, dihedral offset, is being added to the geometry tools software provided by DTM Corporation. The new method, dihedral offset, uses the dihedral angle between connecting facets to calculate the new offset position of the vertex. This method adjusts for the geometry of the part, making the dihedral offset method more accurate than the normal offset and constant offset methods. As an example, Figure 6 shows how two facets and their common vertex are offset a specific distance, d, using each of the three methods.I I Idihedral offsetnormal offsetconstant offsetFigure 6. A visual comparison of the three offset methods available to compensate of the laser beam width.

A visual comparison of the three offset methods available to compensate of the laser beam width. The solid line is the initial position, the dashed line is the final position, the shaded boxes represent the desired amount of offset for each facet.1663.3 Particle Size and Shape Control distinguishable volume elements per unit volume. For SLS technology, process resolution is proportional to the inverse of the particle diameter cubed. As a result, resolution increases substantially as particle size is reduced. Parts made with smaller particles have improved definition as characterized by the sharpness of edges and corners; visually these parts are more models. The impact of particle· shape is more subtle but, in general, parts built with spherical particles. This is particles have improved definition compared to parts build with process.

The impact of particle· shape is more subtle but, in general, parts built with spherical particles. This is particles have improved definition compared to parts build with process. particularly true for materials which undergo limited densification and melting in For such materials, particle shape is not greatly altered in the SLS process.Process resolution isnumbertheInitial SLS materials - wax (LW 2010), polycarbonate (LPC 3000), and nylon (LN 4010) - had mean volume average particle diameters of approximately 100 J.!m. both the polycarbonate and nylon particles is somewhat irregular. In 1994, a nylon material with a particle size of 50 J.!m was introduced. Parts produced with this material were found to have coarser powders. greatly improved feature resolution and definition than those produced with Currently, a number of candidate material systems with small spherical particles are examined. An example of a part made with such a powder is shown in 7.

An example of a part made with such a powder is shown in 7. parts in the photo are made with coarser particles. While the three parts are made chemistries, they all have a similar extent of densification.shapeother two differentFigure 7. The reduction of particle size increases the resolution of the SLS process. The particle sizedecreases from left to right in the SLS parts shown.1674. CONCLUSIONSSLS part accuracy is being improved through a series of projects. In particular, improvements to the beam delivery system, scale/offset calibration and software, and materials have been made or are in development.

In particular, improvements to the beam delivery system, scale/offset calibration and software, and materials have been made or are in development. These improvements are summarized below:Dynamic scanning errors and scanner repeatability errors can be reduced by an order of magnitude with improved galvanometers.With the new galvanometers, the range of dimensional error is reduced by approximately 25 30% to ± 0.25 mm for nylon and polycarbonate, while the range of error is reduced by 40 50% to a value of ± 0.10 mm for polymer coated metals.Optimization of the scanner parameters is ongoing to improve accuracy, surface finish, and repeatability.The revised scale/offset calibration procedure and dihedral beam offset method provides more accurate measurement and compensation for shrinkage and laser beam width.Materials with controlled particle size and shape can produce parts with resolution and definition approaching that of those made with liquid-based processes.5.

proving Accuracy of Powder Sintering-based SFF Processes by Metal Deposition from Nanoparticle Dispersion N. B. Crane*, J. Wilkes†, E. Sachs†, S. M. Allen‡Reviewed, accepted August 26, 2005 * Sandia National Laboratories, PO Box 5800 MS 0958, Albuquerque, NM 87185-0958 † Department of Mechanical Engineering, MIT, 77 Massachusetts Ave, Cambridge, MA 02139 ‡ Department of Materials Science and Engineering, MIT, 77 Massachusetts Ave, Cambridge, MA 02139ABSTRACTSolid Freeform Fabrication processes such as three-dimensional printing (3DP) and selective laser sintering (SLS) produce porous parts that must be densified. New steel infiltration methods can produce parts of standard alloy compositions with properties comparable to wrought materials. However, the infiltration process introduces dimensional errors due to both shrinkage and creep—particularly at the high temperatures required for steel infiltration.

However, the infiltration process introduces dimensional errors due to both shrinkage and creep—particularly at the high temperatures required for steel infiltration. A post-processing method has been developed to reduce creep and shrinkage of porous metal skeletons. Tests have achieved over 90% reduction in creep and 50% reduction in shrinkage. In this method, metal is deposited into the porous part from a suspension of metallic nanoparticles. These particles densify at low temperatures to reinforce the bonds and reduce stress concentrations that amplify creep deformation in untreated parts. After treatment, the reinforced parts can be densified by infiltration.INTRODUCTIONSolid freeform fabrication (SFF) processes offer great geometrical flexibility in manufacturing while eliminating the time and expense of part specific tooling. These advantages make them very attractive for creating appearance models and prototypes and for direct production in low volumes [1].

These advantages make them very attractive for creating appearance models and prototypes and for direct production in low volumes [1]. However, direct part production requires much greater process capability than appearance models and prototypes. This work addresses the dimensional accuracy limitations of parts prepared by Three-Dimensional Printing (3DP) and Selective Laser Sintering (SLS) so that they might be more suitable for applications demanding high accuracy and good properties. Both 3DP and SLS produce parts at high rates by binding powder together, typically with a polymer [2,3]. After forming the geometry, the parts consist of ~50–100 µm particles bound together.For prototype and production applications, the material properties are improved by sintering the powder to full density or infiltration with a second material. Typically, large metal parts are produced in steel powder that is infiltrated with bronze with a linear shrinkage of ~1–3% [4,5].

Typically, large metal parts are produced in steel powder that is infiltrated with bronze with a linear shrinkage of ~1–3% [4,5]. The resulting bronze-steel composite is inferior to homogenous steel alloys and is not widely accepted by designers. New infiltration methods can produce homogenous final parts of standard ferrous alloys [6,7]. Homogenous steel infiltration produced parts with composition and properties comparable to wrought D2 tool steel (1.5% C, 12% Cr, 1% V, 1% Mo, Bal Fe).At liquid-metal infiltration temperatures, sintering and creep processes are active. Both can reduce the final dimensional accuracy. The increase in infiltration temperatures moving from261bronze to steel infiltration substantially increases shrinkage and creep deformations during infiltration, reducing the dimensional accuracy.Typically parts are supported by refractory powder during infiltration to reduce creep deflections.

However, many shapes are difficult to pack consistently, and inconsistent packing can increase variation in the final geometry. Additionally, shrinkage differences between the supporting powder and the parts can tear or deform the part during sintering and infiltration.This work describes and demonstrates a nanoparticle binder and treatment method that can dramatically reduce the deflection rates and shrinkage during infiltration so that the final part accuracy is improved. Section 2 will describe this technique. Section 3 describes the test methods used to evaluate the benefits of the technique. Section 4 discusses the improvements in dimensional accuracy that were achieved and the conditions for optimal performance. Section 5 considers the mechanisms by which the improvements were obtained.PROPOSED NANOBINDERSintering shrinkage and creep deformation under self-weight both cause dimensional changes during infiltration.

The rate of shrinkage and creep are both strongly dependent on the size of the bonds between the micron-scale skeleton particles. While sintering can increase these bond sizes, the shrinkage is increased and the parts may still creep during sintering. Alternatively, if the bond size is increased from added material, both creep and shrinkage can be decreased. [8]We have used a dispersion of metal nanoparticles to increase the size of the necks between particles and thus reduce the dimensional changes during post-processing. This dispersion could be added either by direct inkjet printing in 3DP or it could be infiltrated into parts produced by 3DP or SLS, powder injection molding, or other processes after the geometry is formed. The dispersed nanoparticles can easily fit through the pores between the ~75 µm particles to reach the particle bonds.After applying the dispersion to the skeleton particles, the solvent can be evaporated to deposit the particles in the skeleton.

The solvent dries last from the high curvature areas such as the contacts between particles, concentrating the nanoparticles in these regions. On heating, these particles will sinter to enlarge the neck size and reduce the stress in the bonds. As nanoparticles sinter at much lower temperature than micron-scale particles [9,10], the nanoparticles can densify while the surrounding skeleton powder sinters very little if any.In the special case of ferromagnetic nanoparticles deposited into skeletons of ferromagnetic materials, the nanoparticles can be concentrated in the contact locations even more effectively by applying a DC magnetic field during drying.

When a magnetic field is applied along the skeleton, there is a gradient in the field at the contacts between the micron-scale particles that draws the nanoparticles into these regions.The particles in the dispersion can be chosen to be compatible with the composition of the final part so that the final material properties are not degraded. In steel parts, iron particles are ideal as adding several percent more iron will have a very small impact on the total part composition. Alternatively, a material with a lower creep rate could be chosen that might create bonds with a higher creep resistance, but the impact on the final part properties must be considered. Any262composition gradients created will exist over a few microns and will be reduced by diffusion at high temperatures.This nanoparticle binder process was tested using a gas-atomized, spherical 410 SS (12 Cr, 0.15 C, Balance Fe) skeleton powder sieved to 63–90 µm size.

The nanoparticle dispersion was a commercial product called NanoIron produced by NanoMat. It consists of monodisperse iron nanoparticles with an average size between 7–10 nm. The solvent is a proprietary organic solvent with a boiling point of 287°C. The typical concentration of the nanoparticles was 0.5 vol%.EXPERIMENTAL METHODSThe nanobinder process was tested using a simple test bar with an unsupported overhang as illustrated in Figure 1. These bars were produced either by 3DP using a traditional polymer binder or by sintering powder in an alumina mold. In both cases they were sintered to 1000°C for 30 minutes to develop handling strength and decompose the polymer binder used in the 3D printed parts.Figure 1 Schematic representation of a deflection test bar in an unsupported position (left) and supported position (right).The nanoiron was applied to the bars by infiltration at room temperature.

The solvent was dried at 150°C and the nanoparticles sintered at 700°C for 10 minutes. In most bars, the application process was then repeated to increase the quantity of deposited iron. The drying was done in a nitrogen atmosphere and the sintering in forming gas (95% Ar, 5% H2) to minimize oxidation of the nanoparticles. After applying the nanoiron, some bars were “presintered” by heating them in a supported position (upside down) to 1100°C or 1300°C and holding 60 minutes. The creep strength of the bars was tested by heating in an unsupported position to 1284°C and measuring the final deflection of the tip.The creep and shrinkage performance of the specimens was measured by heating both nanobinder treated and untreated specimens through the same heat cycle and comparing the creep deflections of the tip. The tip deflections were measured using either an optical comparator or images from a flat bed scanner.

The tip deflections were measured using either an optical comparator or images from a flat bed scanner. The length of the supported section was also measured using an optical comparator to calculate linear sintering shrinkage. The ratio of tip deflection of treated specimens to untreated control specimens was used to compare the relative benefit achieved under different treatment conditions. Smaller values of the deflection ratio indicate more benefit from the nanobinder treatment.To better understand the performance of the process, the drying rate, oxidation state of the nanoparticles, and number of nanobinder applications were varied.

The drying rates consisted of fast (hold 20 minutes at 150°C and heat to 300°C in 20 minutes, 1–2 scfh gas flow), moderate (hold 20 minutes at 150°C and heat to 300°C in 200 minutes, 1–2 scfh gas flow) and slow (hold 24 hours at 150°C and heat to 300°C over 4 hours, <0.1 scfh gas flow).263The effect of the oxidation state was compared by applying the same amount of nanoiron from the same batch to two skeletons. One portion of the nanoiron was allowed to oxidize in open air for several days before application while the second was maintained in an inert atmosphere. The two bars were deflected at the infiltration temperature.Some bars were placed in a reservoir of nanobinder during drying so that as the solvent evaporated, additional binder would infiltrate into the skeleton.

Using this method, larger quantities of nanobinder could be deposited in a single application and the creep rates could be compared to samples with the same amount of nanobinder added but in multiple applications. The nanoparticles shrink during sintering, potentially breaking the bonds formed between the particles. However, a repeated application of the nanobinder would partially fill the voids.Some samples were also dried under a magnetic field to improve the concentration of nanoparticles at the contact points between particles. A magnetic field was applied along the length of the bar by placing a rare earth magnet at each end of the specimen during the drying phase. The magnets were removed after the 150°C drying phase to avoid heating above their Curie temperature.RESULTSThe nanobinder reduced both the creep deflection (deflection ratio < 1) and the shrinkage.

The quantity of the nanobinder added, oxidation state, drying rate, and applied magnetic fields each had a significant impact on the nanobinder performance. The best results were obtained by using unoxidized suspensions dried under a magnetic field, with multiple applications. When drying without a magnetic field, better results were achieved by drying at faster rates. In the best case, nanoiron treatment reduced the deflection by 95% as illustrated in Figure 2. The sintering shrinkage of this specimen was also reduced by 60%. This was achieved by applying 2.7 wt% nanoiron in four application cycles dried under a magnetic field. The effect of each of these variables is discussed below.Figure 2 Photo comparing the deflection of 410 stainless steel bars with and without nanoiron treatment. The treated bar received four applications and was dried under a magnetic field.

The treated bar received four applications and was dried under a magnetic field. Both bars were deflected without prior sintering except the 1000°C sintering required to create the initial geometry.264DISCUSSION OF RESULTSDrying RateFigure 3 compares the deflection ratios of bars dried under different conditions as a function of the quantity of nanobinder added to the skeletons. In general, the tip deflection decreases with increasing nanobinder deposition as would be expected. However, greater benefit is obtained at moderate and quick drying rates than at slow drying rates. This is due to changes in the distribution of the nanoiron deposits with drying speed. At higher drying rates, the deposits are concentrated at the outside of the skeleton where the bending stresses are highest. Thus, greater improvement is obtained than if the deposit were distributed more uniformly.

Thus, greater improvement is obtained than if the deposit were distributed more uniformly. The deposit distribution was observed by drying nanoiron in porous cylindrical copper specimens at fast, moderate, and slow drying rates. The parts were sectioned, and EDX spectroscopy was used to measure the relative iron deposit level as a function of distance from the free drying surface. The EDX measurements were made sampling the entire field of view in the SEM at a magnification of 500X. Figure 4 shows that the faster drying rates deposit more material at the outer edge. Even at slow drying rates, the concentration is somewhat higher at the exterior.

Figure 4 shows that the faster drying rates deposit more material at the outer edge. Even at slow drying rates, the concentration is somewhat higher at the exterior. Previous studies [11,12] have shown that the drying conditions can be adjusted to concentrate the deposits on the interior or exterior or uniformly throughout.10.9n o i t c e l f e D0.80.70.6lo r t n o C0.50.4/ t s e T0.30.2Quick DryModerate Dry0.1Slow Dry0slow dry Multiple App0.00% 0.20% 0.40% 0.60% 0.80% 1.00% 1.20%Weight Fraction NmFe AddedFigure 3 Summary of deflection data as a function of drying conditions and quantity of binder material added to the part. Deflection decreases with faster drying rates, increased number of applications, and increasing quantity of deposited material.Multiple Binder ApplicationsOn heating the nanoparticle deposits, they sinter and shrink. The shrinkage can break the bonds between particles as seen in Figure 5.

The shrinkage can break the bonds between particles as seen in Figure 5. However, if the nanobinder is reapplied the cracks can be repaired and on heating, any new cracks will be smaller than before as illustrated. Indeed, when varying amounts of nanoiron were added in a single application, the deflection ratio was approximately 0.85 and did not decrease at larger quantities. However, when the nanoiron was applied in multiple applications, the deflection ratio decreased and showed increasing benefit with increasing deposit volume as shown in Figure 6.265Dried NanoIron distribution. c n o C n o r Id e r u s a e M)%tw(14 12 10 8 6 4 2 0Moderate Rate Quick Dry Slow Dry Untreated02468Distance From Top (mm)Figure 4 Distribution of iron deposits in copper plugs for various drying conditions. Concentration measurements made using EDS across the whole field of view at a magnification of 500X down the centerline of a cylinder of sintered copper powder.

Concentration measurements made using EDS across the whole field of view at a magnification of 500X down the centerline of a cylinder of sintered copper powder. The faster drying rates concentrate more material in a thinner region on the outside of the sample. Since the bending stresses are highest on the outside, the effectiveness of the nanobinder increases.Figure 5 Applying the nanobinder multiple times both enlarges the bonds and heals cracks formed by sintering shrinkage. (Top) Schematic illustration of bond repair by multiple applications of nanobinder. (Bottom) Images comparing bonds sintered to 700°C in samples with 1, 2, and 4 applications and dried under a magnetic field. The size of the cracks in the bonds decreases with additional applications.266Drying in a Magnetic FieldDrying under a DC magnetic field dramatically reduces the deflection ratio compared to samples dried without a magnetic field at a moderate rate as seen in Figure 6.

SEM images of broken interparticle bonds show that the bonds are much larger when dried under a magnetic field than without. This suggests that without the magnetic field, many of the nanoparticles are deposited on the surface of the micron-scale skeleton particles where they do not directly reduce the shrinkage rate.0.40.35Magnet Dry No Magneto i t a R n o i t c e l f e D0.30.250.20.150.10.05a)0b) c)0.00% 0.50% 1.00% 1.50% 2.00% 2.50% 3.00%% Nanoiron AddedFigure 6 Deflection ratios for nanoiron specimens dried with and without a magnetic field. Samples dried under a magnetic field have lower deflections than comparable samples dried without a magnetic field.

Samples dried under a magnetic field have lower deflections than comparable samples dried without a magnetic field. Data includes samples that were presintered at 1300°C and samples deflected directly after processing the nanoiron.Oxidation of the NanoparticlesThe oxidized nanoparticles reduced the creep deflection by 48%, but the same weight of unoxidized material reduces the creep deflection by 64% relative to the untreated case. This difference could be the result of changes in the volume and carbon content of the deposits with oxidation. Iron oxide crystals contain approximately half the iron per unit volume of metallic iron. When the iron oxide is reduced to iron on heating, the deposit density is reduced and the sintering shrinkage for densification increases. Thus, the bonds are more likely to break and more applications of the nanobinder may be required to heal them.The oxides are reduced to iron when the nanoiron deposits are heated, even in an inert atmosphere.

Thus, the oxide must be reduced by the carbon residues of the ligands that stabilize the nanoparticles in the suspension. The carbon-oxygen reaction substantially reduces the carbon content in the skeletons as seen by the difference in carbon content between oxidized and unoxidized nanoiron deposits after heating. Oxidized and unoxidized nanoiron deposits were both heated to 700°C in forming gas and the carbon content of each was measured by combustion infrared detection. The oxidized suspension contained less than 1% carbon compared to 9.7% for unoxidized suspensions under the same conditions. As discussed below, the carbon deposits seem to play an important role in decreasing the creep deflection.Despite these limitations, the oxidized nanoparticles did significantly reduce the creep deformation of the bars. Oxidized nanoparticles are easier to work with than metallic nanoparticles due to their stability in air.

Oxidized nanoparticles are easier to work with than metallic nanoparticles due to their stability in air. Perhaps if carbon can be added to offset that which reduces the iron oxides, the oxidized particles could be effectively used in the nanobinder at lower cost than for metallic particles.267Shrinkage Reduction from the NanobinderFigure 7 shows that the nanobinder consistently reduced the sintering shrinkage of the skeletons. The reduction varies from 40–60% with some increase in reduction with larger deposits. While the creep deflections were sensitive to drying conditions and oxidation state of the suspension, neither of these variables had a significant effect on the shrinkage.

While the creep deflections were sensitive to drying conditions and oxidation state of the suspension, neither of these variables had a significant effect on the shrinkage. This relationship warrants further investigation.3.0001300 C Presinter2.500Oxidized 1300 C Presinter NanoIron)%(e g a k n i r h S r a e n Li2.0001.5001.000No PresinterDried without an applied magnetic field0.5000.0000.00%0.50%1.00%1.50%2.00%2.50%3.00%% NanoIron AddedFigure 7 Variation in shrinkage reduction with quantity of nanoiron added. Total shrinkage reduction is approximately 50% with some increase in shrinkage reduction with increasing nanoiron quantity of deposits. There is not sufficient data to draw conclusions regarding the nature of the relationship. The oxidized nanoparticles are almost as effective at reducing the shrinkage as the unoxidized nanoparticles.

The oxidized nanoparticles are almost as effective at reducing the shrinkage as the unoxidized nanoparticles. Also, the application of a magnetic field during drying does not appear to change the drying shrinkage significantly.MECHANISMS OF CREEP REDUCTIONAdditional experiments were conducted to understand the mechanisms by which the nanobinder treatment reduced the creep deformations of the test specimens. Two important mechanisms were identified: increased bond size and carbon deposition. Each is discussed below.Bond Size IncreasesThe sizes of sintering bonds in parts with and without nanobinder treatment were compared by measuring the size of broken bonds in SEM images of fracture surfaces. Of the two pairs, C-25 & C-18 were presintered at 1300°C while C-29 and C-32 were not. The treated sample in both pairs was dried under a magnetic field to concentrate the deposits in the bonds and improve the deposit uniformity.

The treated sample in both pairs was dried under a magnetic field to concentrate the deposits in the bonds and improve the deposit uniformity. The average bond size and the number of bonds of both treated samples increased relative to the untreated samples. The difference was particularly distinct in the bars that were not presintered as illustrated in Figure 8.268200 µm200 µmFigure 8 Change in bond area with and without nanobinder treatment. After heating the bars through a deflection cycle, the bars were fractured and the area of exposed bonds compared. The nanobinder treated part (left) has much more bond area than the untreated bar (right).Carbon DepositionThe carbon composition of dried nanoiron deposits was measured by combustion infrared detection as 9.7% carbon. While carbon can lower the activation energy for self-diffusion of iron, it also can lower the probability of a successful jump of an iron-atom due to its affinity for vacancy sites.

While carbon can lower the activation energy for self-diffusion of iron, it also can lower the probability of a successful jump of an iron-atom due to its affinity for vacancy sites. Thus, adding carbon can directly decrease iron diffusivity in the alloy at high temperatures. [13] Adding carbon can also change the diffusivity by changing the stable phases present. The pseudo-binary phase diagram of Figure 9 shows that adding carbon to the base 410 SS composition can shift the equilibrium out of the two-phase α-γ region, reducing the grain boundary area available for diffusion and eliminating the higher-diffusivity α phase.1500Liquid)C(1400αγ+αγ+Le r u t a r e p m e T130012001100γγ+CInfiltration Temperature Range10009000.005 C0.01 C0.02410 SS CompositionWeight Fraction CarbonFigure 9 Pseudobinary phase diagram for 410 SS. The basic 410 composition with 0.16% carbon is single phase austenite (γ) between approximately 950°C and 1200°C.

The basic 410 composition with 0.16% carbon is single phase austenite (γ) between approximately 950°C and 1200°C. If the carbon is increased to 0.25%, only the austenite phase is stable between 1000°C and the infiltration temperature. Diffusional creep may be lower in the single-phase region.269The effect of carbon on the deflection of the specimens was tested in two ways. First, a test specimen was infiltrated with a commercial 3DP binder obtained from ProMetal called “Binder S”. This binder leaves residual carbon on decomposition. A second specimen was molded from 420 SS (13 Cr, 0.4 C, Bal Fe) and sintered to 1000°C. Both of these specimens were deflected and compared to a molded 410 SS skeleton prepared in an alumina mold as before and deflected. The carbon content of both samples was measured by combustion infrared detection. Figure 10 shows the deflection ratio of these specimens as a function of the carbon composition. The added carbon clearly reduces the deflection.

Figure 10 shows the deflection ratio of these specimens as a function of the carbon composition. The added carbon clearly reduces the deflection. This suggests that the carbon deposited with the nanobinder is largely responsible for the reduced creep deformation.To test whether the added iron is important, the carbon composition of a number of nanobinder treated skeletons was measured and plotted on the same graph. Magnet-dried samples were used to assure that the deposits were uniformly distributed. These samples show that there is clearly more benefit from samples with both iron and carbon deposited rather than just carbon.The effect of the nanobinder on the deflection performance is particularly dramatic when the samples are not sintered at 1300°C before deflection. When the presinter step was removed from the high-carbon 420 SS parts, the deflection was comparable to a 410 SS part without presinter (deflection ratio ~ 1).

When the presinter step was removed from the high-carbon 420 SS parts, the deflection was comparable to a 410 SS part without presinter (deflection ratio ~ 1). In contrast, the nanoiron treated parts without a presinter step still showed very large improvements relative to similarly processed 410 SS parts (small deflection ratios).1Untreated PartsBinder S Treatedo i t a R n o i t c e l f e D0.90.80.70.60.50.40.3No 1300 C Presinter420 SS - Presinter420 SS - No PresinterNanoBinder Treated0.20.1000.10.20.3NanoBinder - Surface Carbon Nanobinder, Magnet DryIncrease in Carbon Content (wt %)Figure 10 Deflection ratios for treated and untreated bars plotted against the change in carbon composition of the specimen relative to the basic 410 SS composition. Samples include 420 SS material, polymer binder in 410 SS, and nanoiron treatment of 410 SS powder.

Samples include 420 SS material, polymer binder in 410 SS, and nanoiron treatment of 410 SS powder. The carbon treated parts show significant creep reductions, but smaller than the nanoiron treated parts. Additionally, the nanoiron parts follow similar trends with and without a presinter at 1300°C while the 420 SS bars with carbon added have very high deflection ratios (low benefit of carbon) unless it is presintered at 1300°C.This case is particularly important in practical applications because in general it is impossible to presinter the parts at high temperature without incurring creep deformations and/or difficulties with packing the parts in supporting powder. Additionally, the extra step increases the time and cost of the post-processing. It is therefore most desirable to heat the samples for infiltration directly after nanobinder treatment.

Additionally, the extra step increases the time and cost of the post-processing. It is therefore most desirable to heat the samples for infiltration directly after nanobinder treatment. It is in this case when the iron added to the skeleton from the iron nanobinder is most valuable.270These results are consistent with the conclusion that the added carbon reduces the self-diffusion rate in the alloy. If the diffusion rate is reduced, both creep rate and sintering rate are reduced. As a result the sinter bonds are smaller when the carbon content is higher. Without the presintering step, the sintering bonds are very small. Under the high resulting stresses, they quickly fail due to high creep strain.

Without the presintering step, the sintering bonds are very small. Under the high resulting stresses, they quickly fail due to high creep strain. When the bars are treated with nanoiron, the deposits: (1) increase the carbon content to reduce the creep rate and (2) enlarge the bonds between the particles so that the stress is reduced.CONCLUSIONSSteel infiltration processes can produce homogenous, dense material from porous metal skeletons such as those produced by powder injection molding, 3D Printing (3DP), and Selective Laser Sintering (SLS). Despite the improved material properties possible through steel infiltration, at the high infiltration temperatures, dimensional accuracy is reduced by creep deformations under the structure self-weight and sintering shrinkage.Iron nanoparticles have been applied to porous metal skeletons by infiltrating the skeletons with dispersions of the particles and drying the solvent to deposit the particles.

On heating, the nanoparticles sinter quickly to strengthen the bonds between the micron-scale skeleton particles and reducing the creep deformations. The particles have been shown to reduce the creep deflections up to 95% and the sintering shrinkage by up to 60%. The maximum benefit is achieved by repeated application of the nanobinder to heal cracks that form when the nanoparticles shrink. Drying the suspension in a magnetic field concentrates the deposits in the interparticle necks where they strengthen the bonds between the particles. Alternatively, the solvent can be dried quickly so that the particle deposits are concentrated near the outer surfaces of the skeleton where the bending stresses are highest. The benefits are highest for parts treated with unoxidized nanoparticles although iron oxide particles also provide a significant benefit.

The benefits are highest for parts treated with unoxidized nanoparticles although iron oxide particles also provide a significant benefit. Repeated applications of the nanobinder heal cracks formed by sintering shrinkage and improve the performance of the nanobinder.The size of the interparticle bonds is increased significantly by nanobinder deposition. However, the increase in bond size is not sufficient to completely explain the reduced creep deformation achieved by nanoparticle treatment. Carbon deposited with the iron nanoparticles is shown to play a critical role in reducing the creep deflection as well, by reducing the self-diffusivity of iron in the alloy. Although the creep deformation can be reduced significantly by simply adding carbon to the skeleton powder, this approach only works when the parts are sintered at high temperature before heating for infiltration. In contrast, nanobinder treatment significantly reduced the creep deformation even without the sintering step.

In contrast, nanobinder treatment significantly reduced the creep deformation even without the sintering step. Nanobinder treatment was also more effective than adding carbon alone.Iron particles were shown to be more effective at creep reduction than iron-oxide particles. However, the difference may be related to the low carbon content of oxide deposits. The performance of oxidized particles might be improved by depositing extra carbon to compensate for reactions with oxygen.

The performance of oxidized particles might be improved by depositing extra carbon to compensate for reactions with oxygen. Oxidized particles are preferred because they cost less, are not sensitive to air exposure, and may be produced at higher concentrations.271REFERENCES1 Sachs, E., Cima, M., Williams, P., Brancazio, D., and Cornie, J., “Three Dimensional Printing: Rapid Tooling and Prototypes Directly from a CAD Model”, Journal of Engineering for Industry, Vol 114, November 1992, p 481-488.2 Sachs, E., Cima, M., Williams, P., Brancaziu, D., and Cornie, J., “Three DimensionalPrinting Techniques”, US Patent # 5,204,055, April 20, 1993.3 Chua, C. K., Leong, K. F., and Lim, C. S., Rapid Prototyping: Principles and Applictions, World Scientific, 2nd Edition, New Jersey, 2003, pp 175,185.4 Allen, S., and Sachs, E., "Three-Dimensional Printing of Metal Parts for Tooling andOther Applications", Metals and Materials, Vol 6, No.

PROVING ENERGY EFFICIENCY OF CAR CLIMATE CONTROL WITH SLSM. Vetterli†, R. Schmid‡, M. Schmid‡, S. Harke#, T. Durand*, and K. Wegener†† Department of Mechanical and Process Engineering, ETHZ, Zürich 8092, Switzerland ‡ inspire AG, Sankt-Gallen 9014, Switzerland # WeidPlas GmbH, Rapperswil 8640, Switzerland * Automotive Powertrain Technologies Laboratory EMPA, Dübendorf 8600, SwitzerlandAbstractIn Europe, passenger cars are responsible for 12% of CO2 emissions. The European Commission established new regulations to drastically reduce the emissions from 130g to 95g CO2 per km between 2015 and 2021. While the automobile industry is looking at different ways to meet those criteria, the presented industry-driven project aims at reducing energy consumption by up to 30% of air-conditioning (AC) in passenger cars with the introduction of a novel system. The current systems reduce the fuel economy to up to 20% for gas motors and even more for electric cars.

The current systems reduce the fuel economy to up to 20% for gas motors and even more for electric cars. Through Selective Laser Sintering (SLS) design freedom and short production cycles; the design of the AC casing was drastically optimized to increase its contact area with incoming air. To further increase the heat exchange throughout the system, the thermal conductivity of SLS material was improved by incorporation of mineral fillers. The successful implementation of both optimizations led to a CO2 emission reduction of around 50% for the climate control of passenger cars.IntroductionThe emission of greenhouse gases into the atmosphere is believed to be one of the key reasons for the steady increase of temperature throughout the globe provoking an increase in climatic catastrophes. In 2012 within the European Union, the overall emission of carbon dioxide, the main greenhouse gas, attained a value of 3.8 billion tonnes from which 12% originated from passenger cars [1].

In 2012 within the European Union, the overall emission of carbon dioxide, the main greenhouse gas, attained a value of 3.8 billion tonnes from which 12% originated from passenger cars [1]. To tackle this problem, the European union has set emission reduction targets of 130g CO2/km by 2015, lowering even further to 95 g CO2/km in 2021 [2]. The automotive industry investigates different ways of increasing the energy efficiency of their fleet like weight reduction, optimization of combustion engine or develops viable electric/hybrid cars concepts. Farrington [3] showed that using air conditioning could reduce the fuel economy of mid-size passenger cars by 20% and even limited the autonomy range of an electric Mitsubishi iMiev by 46% [4].

Farrington [3] showed that using air conditioning could reduce the fuel economy of mid-size passenger cars by 20% and even limited the autonomy range of an electric Mitsubishi iMiev by 46% [4]. The potential energy savings are important when improving AC in passenger cars; the herein presented project combines thermodynamics with SLS to achieve this goal.Conventional air conditioning systems take outside air in at the front window before being cooled down by the AC system and flood the passenger cabin. During summer, the sun radiation on the hood leads to increase in temperature of the air taken in as showed in Figure 1. This temperature difference decreases with increasing driving speed, however for "STOP and GO" situations, especially in cities, the energy increase of incoming air is important and directly affects the amount of energy needed by the AC to cool down the air taken in.

The narrower the air intake, the less area will be affected by the radiations and the less the temperature difference will be. The current project aims at using cool air from the passenger cabin to counter the uptakeof additional energy during suction before being brought to the AC evaporator and thus diminish the AC energy consumption as simulated in Figure 1.Figure 1 Left: Effect of air intake geometry on temperature difference between ambient outside air and air at partition wall depending on the driving speed adapted from Grossmann [5]; Right: Graphical representation of box configuration for pre-conditioning of air taken in at the front windshield.The new AC concept is composed of an outer and an inner box as shown in Figure 1. The incoming air is suctioned through a narrow air intake at the front window, directed to the inner box and finally to the AC condensator.

The incoming air is suctioned through a narrow air intake at the front window, directed to the inner box and finally to the AC condensator. At the same time, recycled air from the passenger cabin, with comparatively much lower energy content, is expelled through the outer box surrounding the inner box. The inner box wall acts as "heat exchanger" and must therefore conduct heat well enough to optimize the exchange between the air fluxes. In comparison to the “air recirculation” option in cars, this novel system would hinder the need of removing of odours/exhalations of passengers as well as controlling humidity levels [6].With help of design freedom enabled by SLS as well as its rapid implementation, different designs for the AC box are investigated to optimize its configuration for improving heat exchange between incoming and outgoing streams.

The conventional injection moulded material to produce air ducts and air-water separation units for the climate control system used by the industrial partner is a talcum-filled polypropylene (PP). Therefore, the thermal properties of the SLS material should present similar values to this widely used material. For this purpose and due to the rather limited portfolio of available SLS materials, the development of potential novel SLS materials based on a novel SLS polyolefin, iCoPP presented by Schmid [7], is investigated. The resulting AC box will be tested in real conditions in a premium car and compared to conventional AC systems during standardized driving cycles. The tests will be conducted in a climate chamber rendering the conditions of a sunny summer day of Western Europe.MethodsSLS Materials DevelopmentThe automobile industry uses polypropylene (PP) in combination with mineral fillers to increase different material properties.

The injection moulding polypropylene of the industrial partner is filled with talcum and presents a thermal conductivity of 0.35 W/mK. Due to the low thermal conductivity of polyamide 12 (0.2 W/mK), the most used polymer for SLS, it will not be taken into account. To obtain an SLS material presenting such thermal conductivity, differentmineral fillers have been considered for reinforcing iCoPP, a novel polyolefin suitable for the SLS process. The iCoPP is a polypropylene copolymer and currently the only SLS polymer on the market to present a combination of high elongation at break (~200%) and reasonable stiffness (800-1000 MPa). It was dry blended with the fillers to increase its thermal conductivity. Moreover, some other promising SLS materials available in the market were investigated and compared with the in-house developments. Table 1 summarizes the materials investigated for this project. The material TPP is a talcum reinforced PP produced by extrusion.

Table 1 summarizes the materials investigated for this project. The material TPP is a talcum reinforced PP produced by extrusion. Talcum is widely used to increase thermal properties in polymers. For this reason the optimal type of talcum to use, fine or coarse, for the SLS process was determined in a pre-screening round with fine Luzenac, with a D95 diameter of 7.7 µm, and Luzenac 1445, D95 of 49.2 µm. The impact on thermal properties and flowing behaviour of both types with iCoPP were determined.

The finer talcum proved to worsen the flowability so much that it could not be measured on the RPA as well as shrink even further the sintering window; only the coarser talcum could thus be considered.Table 1 List of the assessed materials during the project.ObtainedPolymerTypeCommentsA1polypropylene copolymer, D50,vol 47.3 µmDevelopmentCommercialiCoPPTPPMicrofol SinterplastDuraform HSTA2A3BCD20wt% calcium carbonate (CaCO3) surface treated (Omyafilm 753 – FL, Supplier: Omya, with D50,vol 3.2 µm) 20wt% talcum (Luzenac 1445,Supplier: Imerys, with D50,vol 15.6 µm) polypropylene reinforced by talcum, obtained by coextrusion, D50,vol 46.2 µm polypropylene reinforced by glass beads, dry blend, D50,vol 68.1 µm Polyamide 12 reinforced by mineral fibers, dry blend, D50,vol 60.5 µmThe characterization of a material suitable for SLS must not only take into account the powder properties but properties of the processed material as well; for the project the methods used to determine the powder and the processed material are presented in Table 2.

The flowability of a powder is essential to help predict its processability while the DSC gives indications about the sintering window. The sintering window is the temperature range in which the semi-crystalline polymer is molten and has not yet begun its crystallization as detailed by Schmid [8]. This value is calculated by the difference between the onset melting and crystallization temperatures obtained during the DSC experiment ((Tm-Tc)onset).

The presence of foreign bodies inside the molten polymer leads to heterogeneous nucleation due to the diminution of the free surface energy at the interface filler-polymer and starts the crystallization at higher temperature [9].Table 2 Characterization methods implemented during the projectStateName Revolution Powder Analyser (RPA)Method descriptionBacklit rotating drum to investigate the powder flowing behaviour [10]Measurement Conditions25 ml powder, counting 128 avalanches at a rotation speed of 0.6 rpmPowderDifferential Scanning Calorimetry (DSC)Determination of thermal and heat transition behaviour during heating and cooling1st heating: 25 - 200°C, 10°C/min Cooling: 200-25°C, -10°C 2nd heating: 25-200°C, 10°C Nitrogen flow 50 ml/min.ProcessedHeat ConductivityHeat Deflection Temperature (HDT)Modulated DSC determines the heat conductivity through a solid body Temperature at which a material shows a certain deformation under a certain loadASTM E1952-11DIN EN ISO 75-2:2013-08/A (1.8 MPa).The heat deflection temperature (HDT) is the temperature at which a slab, loaded with a defined charge, undergoes a given deformation.

The HDT value of the industrial PP lies around 80°C. However, because the AC casing is not mechanically solicited and the comparatively short duration of the tests to a whole lifecycle of car components, the HDT values of the investigated materials should at least withstand temperatures of 50°C. The commercial materials (A1, C, and D) and the TPP (B) were produced on a DTM Sinterstation 2000 with optimized process parameters. The development materials A2 and A3 were extruded on a laboratory extruder at a temperature of 150°C using a rectangular die with an opening of 10x8 mm. This extrusion process was used in order to obtain a homogeneous distribution of the fillers inside the composites and therefore solely determine their effect on the thermal conductivity of iCoPP composites.Design OptimizationThe AC box was subject to design optimization in order to further increase the heat exchange between the air fluxes. Its design was modified using the software Solidworks and MAGICS.

Its design was modified using the software Solidworks and MAGICS. It was subject to certain restrictions to achieve its successful implementation under the car’s hood. The dimensions had to show an optimal balance between the available place under the hood and the minimal volume to avoid a pressure drop within the aeration system. Secondly, there are only few SLS machines able to produce high volume parts and these machines generally process only the conventional PA12. This was the reason why multimaterial designs were investigated to optimize not only the volume and structure of the box but as well the heat exchange part. Finally the inner box should present the maximum contact area between the incoming and out coming air fluxes.

Finally the inner box should present the maximum contact area between the incoming and out coming air fluxes. Those optimization steps will be assessed by mounting the boxes in the car and measuring the fuel consumption and compressor voltage under real driving conditions.Test Cycles MeasurementsThe energy consumption tests were conducted in real conditions with a Mercedes C-class during which the following parameters were measured: the fuel consumption, the temperature at partition wall as well as the compressor voltage. They took place in the climatic chamber at theEMPA in Dübendorf that reproduced a summer sunny day in Western Europe with a temperature of 30°C and a relative humidity of 65%. The front window of the car was irradiated with halogen lamps at a power of 1000 W/m2 to mimic the solar radiation.

The front window of the car was irradiated with halogen lamps at a power of 1000 W/m2 to mimic the solar radiation. Three different standardized driving test cycles were run (ECE15 1970, NEFZ 1990 and WLTP 2017).The "Worldwide Harmonized Light Duty Test Procedure" is the closest to real driving conditions [11]. The successful materials were processed on an SLS machine in the optimized design of the AC box and compared to reference AC configurations to determine the decrease in energy consumption.

The powder properties show a clear dependency on the type of filler used. A powder is considered as flowing homogeneously and in a repeatable manner when the mean avalanche angle and surface fractal are low and show narrow standard deviations [10]. The way the powder flows is affected by the filler form, spherical fillers like for CaCO3 (type A2) or glass beads for type C hinder less the movement of particles relative to each other than lamellar structures like talcum (type A3) or needles in the case of D. This fact is here confirmed by the avalanche angles as well as the surface fractals of all the investigated powders. Here, the best flowing material is the type B, the particles are not only ellipsoidal but their increased weight compared to pure PP help stabilizing the flowing behaviour.

Here, the best flowing material is the type B, the particles are not only ellipsoidal but their increased weight compared to pure PP help stabilizing the flowing behaviour. The homogeneity of the powder particles is a key factor as can be observed in Figure 2 and it can be observed that fillers inside the particles will be less likely to hinder the powder flow than dry blended fillers. The width of the sintering window is a very sensible parameter and good comparison basis for the processing of the material on the machine. For the developed materials and depending on the morphology of the filler, the sintering window is either not affected, as for type A2, or diminished due to the nucleating effect of the filler, as for type A3.

The type B presents the widest sintering window and fulfils thus the most important characteristics of an SLS powder.Figure 2 SEM pictures at 1000 magnification; Left type B, polypropylene extruded with talcum; Right type A3 obtained by dry blending of coarse talcum and iCoPPThe thermal conductivity is highest for the talcum filled materials A3 and B even though they have a significant difference of 0.13 W/mK. The more intimate contact of PP with talcum obtained as well as its homogeneous distribution during the extrusion process in type B leads to a better conduction of heat through the material than for the dry blending case. The other materials present all better conductivity than plain polyamide but are still half as good conductors as the type B. Type C and D show the highest HDT values but due to their flowing behaviour coupled with thermal conductivity, they are less interesting for SLS processing than type B.

Type C and D show the highest HDT values but due to their flowing behaviour coupled with thermal conductivity, they are less interesting for SLS processing than type B. Type B with its good powder properties combined to high heat conductivity is definitely the material of choice for the casing of the AC box.Design OptimizationFigure 3 Optimization steps of AC box designs; Left: First step; Middle: Increased surface area (2x); Right: PA12 structure and insert out of developed material (area 2.5-3x)Figure 3 presents the evolution of the box design from basic box (left) to a more complex and combining multi-material box. For the second iteration, only the inner box was optimized to double its surface area. However, due to the big volume of the inner box, there exists few machines on the market able to process such dimensions and the ones having a large enough building platform generally only process PA12.

The final version of the inner box is a structured body produced with PA12 and high surface area inserts. These pull-outs can be produced using the different SLS materials designed for and investigated in this work. Thus, the novel AC system is mainly composed of PA12 and, at the positions of desired maximum heat exchange (honeycomb structures in Figure 3 - right picture) with high conductivity optimized material. The outgoing cool cabin air flows around the inner box and heat exchange happens at the wall with the outside hot incoming air.

The outgoing cool cabin air flows around the inner box and heat exchange happens at the wall with the outside hot incoming air. The following WeidPlas inner box configurations are tested in real conditions:WP1: structure and inserts produced with PA12, tests are performed with and without evaporative cooling (EC)WP2: structure produced with PA12, inserts made out of material type A with EC - WP3: structure produced with PA12, inserts made out material type B with ECThe evaporative cooling process enables to further cool down the cabin outflow by using water’s high enthalpy of vaporization. The cool airflow from the cabin passes through a curtain of sprayed water, obtained by condensation of the hot outside air.

The cool airflow from the cabin passes through a curtain of sprayed water, obtained by condensation of the hot outside air. By flowing through the water curtain, the cool air will lower its temperature while increasing its humidity.Test Cycles MeasurementsFigure 4 Air intake temperature at partition wall during the WLTP test cycles for the configurations Premium Vehicle 1 and 2 (PV1-2), the Mercedes C-Class configuration (Serie C-Class) and the developed boxes (WP1-2-3).Figure 4 presents the temperature evolution of intake air at the partition wall during the WLTP driving test with the different AC configurations. The temperature ranges of the three Weidplas concepts (WP1-2-3) present clearly reduced temperature of the air transmitted to the AC evaporator, needing thus less energy to cool the air before flooding the passenger cabin.

The behaviour presented in Figure 1 is confirmed; the prepared air with the new concept is much cooler than the conventional systems, with either broad or narrow intakes. The effect of material conductivity is clearly observed with the lowest temperature profile for the type B, the most thermally conductive SLS material. The temperature decreases with increasing thermal conductivity of the AC box; it confirms the original idea of increasing heat transfer between two fluids through a wall to cool down the air sucked into the car.An important feature that needs to be mentioned is the surface roughness of the SLS inserts. The intrinsic roughness of SLS parts is here a decisive point for increasing the convective heat transfer at the wall interface between the two air fluxes.

The intrinsic roughness of SLS parts is here a decisive point for increasing the convective heat transfer at the wall interface between the two air fluxes. This effect will need to be taken into account when the AC box is produced via injection moulding where a smooth surface is standard.Fuel Consumption of AC SystemFigure 5 Fuel consumption of air conditioning system and air temperature at partition wall for the WLTP cycle at 30°C and 65% relative humidity.Figure 5 shows the results this project aimed for; it gives the fuel consumption of the AC system in relation to the air temperature at partition wall for the different AC configurations. Just like in the case of Figure 4, the Weidplas concept box values are close together while the reference systems are scattered at high consumption values. In the optimal case, WP3 - type B, there is a reduction of 50% in fuel consumption compared to the premium vehicle solutions.

In the optimal case, WP3 - type B, there is a reduction of 50% in fuel consumption compared to the premium vehicle solutions. The voltage of the refrigerating compressor diminished from 7.2 V in the case of the Premium Vehicle 1 to 5.2 V for the WP3, a diminution of 28% of the energy consumption of the AC system. However, even the system made entirely out of PA12 shows a drastic amelioration of the AC system on an energetic point of view. The thermal conductivity increase from PA12 (0.2 W/mK) to TPP (0.57 W/mK) enables fuel savings of up to one third. The combination between increased surface area, higher thermal conductivity of the material as well as rough surface allows the overall heat transfer to become more efficient and thus help reducing CO2 emissions. This feature will certainly be of major interest to the automotive industry around the world.ConclusionThe efficiency of air conditioning in private cars has been successfully increased.

This feature will certainly be of major interest to the automotive industry around the world.ConclusionThe efficiency of air conditioning in private cars has been successfully increased. The effect of precooling the intake air at the front window with reused air from the passenger cabin has been found to diminish the fuel consumption of AC system to up to 50%. The novel concept was confirmed using the SLS technology advantages for the inner box design. The construction optimization enabled a drastic surface area increase to enhance the heat exchange between air fluxes while the use of inserts produced out of different thermally conductive SLS materials. To further improve the system, new SLS materials with improved thermal conductivity were identified and characterized before being implemented inside the AC box.

To further improve the system, new SLS materials with improved thermal conductivity were identified and characterized before being implemented inside the AC box. The effect of thermal conductivity was found to give a decisive boost to the fuel savings during standardized test cyclesThis novel AC system enables energy savings for cooling down cars in summer. In Europe, warm sunny days are less frequent than cold, rainy days and this is the reason why this system will be investigated under winter conditions with reuse of warm air from the passenger cabin. This feature is especially important for electric vehicles during winter. Up to now the driving range of electric cars was kept constant to the detriment of the driver’s comfort. Preliminary results tend to confirm an efficiency increase for warming up as well.AcknowledgmentsThe authors gratefully acknowledge the federal Commission for Technology and Innovation of Switzerland for funding the project 14793.1 PFIW-IW.

proving Linear Weld Density in Ultrasonically Consolidated PartsG.D. Janaki Ram, Y. Yang, J. George, C. Robinson and B.E. Stucker Department of Mechanical and Aerospace Engineering, Utah State University Logan, UT 84322-4130, USAReviewed, accepted September 14, 2006AbstractUltrasonic consolidation is a novel additive manufacturing process with immense potential for fabrication of complex shaped three-dimensional metallic objects from metal foils. The proportion of bonded area to unbonded area along the layer interface, termed linear weld density (LWD), is perhaps the most important quality attribute of ultrasonically consolidated parts. Part mechanical properties largely depend on LWD and a high level of LWD must be ensured in parts intended for load-bearing structural applications. It is therefore necessary to understand what factors influence LWD or defect formation and devise methods to enhance bond formation during ultrasonic consolidation.

It is therefore necessary to understand what factors influence LWD or defect formation and devise methods to enhance bond formation during ultrasonic consolidation. The current work examines these issues and proposes strategies to ensure near 100% LWD in ultrasonically consolidated aluminum alloy 3003 parts. The work elucidates the effects of various process parameters on LWD and a qualitative understanding of the effects of process parameters on bond formation during ultrasonic consolidation is presented. The beneficial effects of using elevated substrate temperatures and its implications on overall manufacturing flexibility are discussed. A preliminary understanding of defect morphologies and defect formation is presented, based on which a method (involving surface machining) for minimizing defect incidence during ultrasonic consolidation is proposed and demonstrated. Finally, trade-offs between part quality and build time are discussed.1.

Finally, trade-offs between part quality and build time are discussed.1. IntroductionUltrasonic Consolidation (UC) is a novel additive manufacturing process developed by Solidica Inc., USA, utilizing the principles of ultrasonic welding [1]. The process builds up the rough part shape by ultrasonically welding or consolidating thin metal foils (typically 150 (cid:80)m thick). This ultrasonic addition is combined with 3-axis CNC milling to produce geometric details. The Solidica Form-ation UC machine (Fig.1), commercially introduced by Solidica in 2000, is an integrated machine tool which incorporates an ultrasonic welding head, a foil feeding mechanism, a 3-axis milling machine, and software to automatically generate tool paths for material deposition and machining. Part fabrication takes place on a firmly bolted base plate (typically of the same material as the foil being deposited) on the top of a heat plate.

Part fabrication takes place on a firmly bolted base plate (typically of the same material as the foil being deposited) on the top of a heat plate. The heat plate maintains the substrate at a set temperature allowing the deposition process to be carried out at temperatures ranging from ambient to 350(cid:113)F.Fig.2 illustrates the basic UC additive manufacturing process. In this process a rotating ultrasonic sonotrode travels along the length of a thin metal foil placed over the substrate. The thin foil is held closely in contact with the substrate by applying a normal force via the rotating sonotrode. The sonotrode oscillates transversely to the direction of welding at a frequency of 20 kHz and at a user-set oscillation amplitude, while traveling over the metal foil. The combination of normal and oscillating shear forces results in generation of dynamic interfacial stresses at the interface between the two mating surfaces [1-3].

The combination of normal and oscillating shear forces results in generation of dynamic interfacial stresses at the interface between the two mating surfaces [1-3]. The stresses produce elastic-plastic deformation of surface asperities, which breaks up the oxide film, producing relatively clean metal surfaces692under intimate contact, establishing a metallurgical bond. Oxide films, broken up during the process, are displaced in the vicinity of the interface or along the weld zone. Local temperatures at the interface and the surrounding affected region (about 20 (cid:80)m) can reach up to 50% of the melting point of the material being deposited [3]. After depositing a strip of foil, another foil is deposited adjacent to it. This process repeats until a complete layer is placed. After placing a layer, a computer controlled milling head shapes the layer to its slice contour. This milling can occur after each layer or, for certain geometries, after several layers have been deposited.

This milling can occur after each layer or, for certain geometries, after several layers have been deposited. Once the layer is shaped to its contour, the chips are blown away using compressed air and foil deposition starts for the next layer.Fig.1. Solidica Form-ation ultrasonic consolidation machine.Fig.2. Schematic of the ultrasonic consolidation process.693Ultrasonically consolidated parts typically show metal-to-metal bonded regions and a few defects (oxide accumulated regions and/or physical discontinuities) along the layer interfaces, as noted by Kong et al. in their studies on UC of Al alloys [4,5]. A parameter called “linear weld density” (LWD) is generally used to represent the proportion of bonded area in relation to the total interface length [4], and is the most important quality attribute of ultrasonically consolidated parts. Mechanical properties of ultrasonically consolidated parts largely depend on %LWD between metal foils.

Mechanical properties of ultrasonically consolidated parts largely depend on %LWD between metal foils. It is therefore necessary to minimize defects and ensure a high level of LWD in ultrasonically consolidated parts for use in load-bearing structural applications. In this context, selection of appropriate process parameters plays a key role. In their studies on UC of alloy 3003, Kong et al. [4] have presented a preliminary understanding of the effects of oscillation amplitude, welding speed and normal force on bond formation and identified a general process window for successful part fabrication based on microscopic studies and peel tests. The highest LWD reported in this study was 87%. The authors, however, did not include substrate temperature as a variable and their experiments were conducted at room temperature.The mechanism of bond formation during ultrasonic welding of metals has long been investigated, although considerable debate still exists in this matter [2,3].

According to the current understanding, bond formation during ultrasonic welding of metals involves three mechanisms: i) removal of surface oxide layers, ii) plastic deformation at the interface, and iii) diffusion of metal atoms across the interface. Plastic deformation aids in bond formation in three ways i) it helps break the surface oxide layers causing metal flow through the cracks generated in the oxide layer by the action of ultrasonic energy and, perhaps more importantly, remove the broken oxide scales away from the bonding region, ii) it helps increase interfacial contact area by causing metal flow into the gaps existing at the interface due to surface roughness, iii) it helps bring nascent metal surfaces in intimate contact across which bonding sets in (diffusion of metal atoms across the interface helps, but bonding can occur without it). Therefore, plastic deformation at the interface, in a way, can be considered the enabling mechanism for the other two mechanisms.

Therefore, plastic deformation at the interface, in a way, can be considered the enabling mechanism for the other two mechanisms. Having noted the importance of plastic deformation, use of elevated substrate temperatures is potentially beneficial in promoting bond formation However, practically no published information is available on the effect of substrate temperature during UC.Another important aspect is the origin of defects that show up at the layer interfaces, on which very little is known. It is important to gain greater insights into this aspect and precisely identify the origin of defects, based on which one can devise suitable techniques to minimize the incidence of defects. This paper describes our ongoing efforts to improve LWD in ultrasonically consolidated parts focusing on i) process parameters effects, including substrate temperature, and optimization, and ii) origin of defects and defect elimination through process improvements.

The trade-offs among part quality, build time and manufacturing flexibility are discussed in the context of proposed strategies for improving LWD in ultrasonically consolidated parts.2. Experimental Work 2.1 MaterialsThe material used for these UC experiments was Al alloy 3003 (nominal composition by wt.%: Al-1.2Mn-0.12Cu) foil, 150 (cid:80)m thick and 25 mm wide, obtained from Solidica, Inc., USA. Deposition experiments were conducted on an Al 3003 base plate (dimensions: 355x355x12 mm) firmly bolted to the heated platen in the machine.6942.2 Parameter optimization experimentsA design of experiments (DOE) approach was adopted to systematically evaluate the effects of process parameters and to identify the optimum parameter combination. The process parameters and the levels selected for evaluation in this study are shown in Table 1. Variation of each parameter at four different levels was considered necessary to assess any non-linear effects.

Variation of each parameter at four different levels was considered necessary to assess any non-linear effects. Specific levels for each of the parameters were selected based on preliminary experiments, and machine setting limits. A Taguchi L16 orthogonal array was utilized in the present study to determine the effects of individual process parameters. Interacting influences between two or more process parameters were not assessed. Table 2 lists all the parameter combinations used for deposition experiments. The experimental runs were randomized and each of the 16 runs was repeated twice. Each run consisted of depositing four layers of foil one over another (Fig.3). The welding direction was reversed for each layer. The ultrasonic oscillation frequency was maintained constant at 20 kHz for all the experimental runs.The response parameter employed in the present study was %LWD.

The ultrasonic oscillation frequency was maintained constant at 20 kHz for all the experimental runs.The response parameter employed in the present study was %LWD. Analysis of Variance (ANOVA) was performed taking LWD as the response parameter, following standard statistical procedures [6]. Following the initial round of experiments, additional experiments were conducted at relatively lower welding speeds (24, 20, 16 and 12 mm/s) keeping all other parameters constant at their optimum levels. These experiments were considered necessary to unambiguously assess the effect of welding speed on LWD.Table 1. Parameters and levels selected for UC experiments.ParameterLevel 1Level 2Level 3Level 4Oscillation amplitude ((cid:80)m) Welding speed (mm/s)1028133216361940Normal Force (N) 1450 75 Temperature ((cid:113)F)1600 1501750 2251900 3002.3 MetallographyAll the deposits were metallographically examined to estimate %LWD.

Samples corresponding to longitudinal and transverse sections were extracted from each of the deposits and were prepared for microstructural study following standard metallographic practices. Studies on both deposit center and end portions were considered necessary to assess any local variations in bond formation. Microstructural observations were conducted using an inverted light microscope (Zeiss Axiovert 100A) on as-polished samples. Pictures were taken at a number of locations on each sample (at least 12 picture frames) and were used to measure the average LWD in each deposit. The interface between the base plate and the first layer was not considered in LWD measurement for reasons described in the next section. The LWD on a picture frame was measured using the following formula:%LWD(cid:32)(cid:170) (cid:171) (cid:172)Bonded Totalinterfaceinterfacelengthlength(cid:186) (cid:117)(cid:187) (cid:188)100695Table 2.

Taguchi L16 experimental matrix and the results of linear weld density measurements.Run #1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16Amplitude ((cid:80)m) 10 16 19 13 13 19 10 16 13 10 19 16 19 16 13 10Speed (mm/s) 28 40 32 36 40 28 32 36 28 36 40 32 36 28 32 40Force (N) 1450 1600 1750 1900 1750 1900 1600 1450 1600 1750 1450 1900 1600 1750 1450 1900Temperature ((cid:113)F) 75 75 75 75 150 150 150 150 225 225 225 225 300 300 300 300LWD (%) Deposit 1 Deposit 2 Average 19 60 65 26 33 74 26 52 58 54 37 59 68 88 62 4017 50 69 24 31 78 22 58 62 46 35 55 72 92 58 4418 55 67 25 32 76 24 55 60 50 36 57 70 90 60 42Fig.3. Schematic of the UC deposits.2.4. Surface machining experimentsBased on metallographic observations, sonotrode-induced surface roughness on the deposited foil surface was identified as the major source of defects in ultrasonically consolidated parts.

This led us to the idea of employing an intermediate surface machining step after depositing each layer and prior to subsequent layer deposition to remove the sonotrode-induced surface roughness on the just deposited layer/foil surface. Experiments were conducted to demonstrate this idea at various welding speeds (at both ambient and elevated substrate temperatures) by removing about 30 microns thick of material (based on observations of the scale of sonotrode-induced surface roughness, which is about 10-15 microns) on the foil surface after depositing each layer using the CNC mill that is part of the Solidica Form-ation machine. Table 3 lists all the parameter combinations used for deposition experiments with surface machining (no replication). The surface machining step was incorporated into the machine code696so that the process proceeds from one step to the next automatically.

The surface machining step was incorporated into the machine code696so that the process proceeds from one step to the next automatically. No modifications to the regular part building sequence utilized by the UC process were made except the introduction of an intermediate surface machining step.Table 3. Welding speeds and substrate temperatures used in deposition experiments with surface machining and corresponding LWD results. Other parameters were kept constant at their optimum levels (oscillation amplitude: 16 (cid:80)m, Normal force: 1750 N). Substrate Temperature ((cid:113)F) 75 300 75 300 75 300 75 300Welding speed (mm/s)LWD (%)100 100 100 100 90 100 80 90283236403. Results and Discussion 3.1 Effect of process parameters 3.1.1 Linear weld densityNo significant differences in LWD were noticed between center and end (start/stop) portions of a deposit, as can be seen, for example, in the microstructures (longitudinal section) of Run # 6 shown in Fig.4.

The dark regions seen along the layer interfaces are the unbonded regions. No specific relationship was found to exist between the occurrence of defects and the welding direction. Similarly, no significant differences in LWD were observed across the deposit width as can be seen, for example, in the microstructures (transverse section) of Run # 8 shown in Fig.5. However, considerable differences in LWD across the deposit width were noticed in higher aspect ratio deposits made in a subsequent study [7].Three distinct types of defect morphologies were observed, as can be seen in Fig.4 and Fig.5: (i) line-like defects, (ii) parabola-like defects, and (iii) point-like defects. While all the three types of defects were present to some extent in all the deposits, certain broad trends were observed. The occurrence of line-like defects was found to be more frequent in samples exhibiting a very low %LWD, deposited using a low level of oscillation amplitude and/or normal force.

The occurrence of line-like defects was found to be more frequent in samples exhibiting a very low %LWD, deposited using a low level of oscillation amplitude and/or normal force. The parabola-like defects were always observed to be pointed downwards with flat tops and curved bottoms. Parabola-like defects were found to be more frequent in samples exhibiting medium or medium to high weld density levels, while samples exhibiting very high weld density levels usually showed only point-like defects. These observations, however, were not statistically verified during this study, and further studies are planned to further understand defect morphologies and correlate their occurrence to bond formation mechanisms and process parameter effects.697(a)(b)Fig.4. Microstructures of Run # 6 deposit (longitudinal section): (a) center, (b) end (start/stop) portion. Horizontal arrow show welding direction for each layer.

Microstructures of Run # 6 deposit (longitudinal section): (a) center, (b) end (start/stop) portion. Horizontal arrow show welding direction for each layer. D1, D2 and D3 show line-like, parabola-like, and point-like defects, respectively.(a)(b)Fig.5. Microstructures of Run # 8 deposit (transverse section): (a) center, (b) edge. D1, D2 and D3 show line-like, parabola-like, and point-like defects, respectively.During ultrasonic welding, the mating surfaces come in contact with each other at surface asperities under the influence of an applied normal force. As noted earlier, the combination of normal and oscillating shear forces results in generation of dynamic interfacial stresses between the two mating surfaces. The stresses produce elastic-plastic deformation of surface contact points, which breaks up the oxide film, producing relatively clean metal surfaces under intimate contact, establishing a metallurgical bond.

However, this process leaves some unbonded regions along the interface. Probable causes for these unbonded areas include: (i) lack of complete698contact between mating surfaces due to surface roughness and/or entrapped air, (ii) persistence of surface oxide layers preventing intimate nascent metal contact, and (iii) accumulation of removed surface oxides at localized regions along the interface. For these reasons, 100% bonding is difficult to achieve in ultrasonic welding. Further, while depositing a layer, sonotrode motion on the foil can result in a very rough surface with hills and valleys, as can be seen on the top layer in Fig. 4a and Fig.5a (the foil stock used in this study has a very fine, mirror-like surface finish). This sonotrode-induced roughness on the deposit surface can prevent effective surface contact during subsequent layer deposition and the regions corresponding to valleys can manifest into defects.

This sonotrode-induced roughness on the deposit surface can prevent effective surface contact during subsequent layer deposition and the regions corresponding to valleys can manifest into defects. This hypothesis is supported by the occurrence of parabola-like defects with flat tops and curved bottoms, whose size closely matches with the roughness scale induced on the foil surface due to sonotrode motion. Therefore, it is our belief that this sonotrode-induced roughness is a major source of defects in ultrasonically consolidated parts.The results of linear weld density measurements are presented in Table 2. Variation in LWD within a deposit and from one deposit to another deposit of any one experimental run was found to be generally less than 15%. Representative microstructures of some of the deposits (longitudinal sections) are shown in Fig.6.

Representative microstructures of some of the deposits (longitudinal sections) are shown in Fig.6. The first layer was found to be bonded to the base plate at a considerably higher LWD than the second, third and forth layers in all the deposits. This is mainly attributable to a combination of the absence of an oxide layer on the base plate and a lower surface roughness (absence of sonotrode-induced roughness), as the machine performs surface milling on the base plate just prior to depositing the first layer. For this reason, the interface between the first layer and the base plate was not considered in LWD measurements.The LWD was found to change considerably among the 16 experimental runs ranging from 18% (Run # 1) to 90% (Run # 14) indicating that bond formation is strongly dependent on process parameters. The results of ANOVA analysis are summarized in Table 4. All the four parameters contributed to statistically significant variations in LWD at a 90% confidence level.

The results of ANOVA analysis are summarized in Table 4. All the four parameters contributed to statistically significant variations in LWD at a 90% confidence level. Oscillation amplitude was found to exert the strongest influence among the four parameters studied. Welding speed and substrate temperature were found to have a similar level of influence on LWD. The normal force was found to have the weakest influence on LWD within the parameter range used in this investigation.Table 4. Results of ANOVA analysis.SourceAmplitude Speed Force Temperature Error TotalSums of squares (SS) 5232 1577 1247 1527 2907 12490Degrees of freedom ((cid:81)) 3 3 3 3 19 31Variance (V) 1744 526 416 509 153F11.40 3.43 2.71 3.32F (table, 3,19) at 90% confidence = 2.40699(a)(b)(c)(d)(e)(f)Fig.6. Microstructures of UC deposits (longitudinal section): (a) Run # 1, (b) Run # 3, (c) Run # 5, (d) Run # 9, (e) Run # 14, (f) Run # 16.

Asummary of observations is presented here.Effect of oscillation amplitude: The LWD increased with increase in oscillation amplitude from 10 to 16 (cid:80)m (Fig.7a). An increase in oscillation amplitude increases the magnitude of oscillating shear forces and, hence, the magnitude of dynamic interfacial stresses at the interface between the two mating surfaces. This would enhance the elastic-plastic deformation at the surface contact points and facilitate easier removal of surface oxide layers and plastic flow. It is likely for these reasons the deposits showed an increase in LWD with increase in oscillation amplitude from 10 to 16 (cid:80)m. Further increase in oscillation amplitude to 19 (cid:80)m resulted in a small drop in LWD (Fig.7a). The reasons for this are unclear, but may be due to one or more of the following reasons.

Further increase in oscillation amplitude to 19 (cid:80)m resulted in a small drop in LWD (Fig.7a). The reasons for this are unclear, but may be due to one or more of the following reasons. When the amplitude is too high, excessive stresses developed at the interface may break up already formed bonds just behind the moving sonotrode, particularly if these higher stresses also resulted in excessive strain hardening, resulting in a lower weld density. Also, at a fixed ultrasonic oscillation frequency, the time available for completing a cycle is fixed. Thus when the amplitude is set at a higher level, the sonotrode must accelerate and decelerate more rapidly in order to complete a cycle, which can affect the sonotrode operational performance.Effect of welding speed: The LWD increased with decrease in welding speed from 40 to 28 mm/s (Fig.7b).

Welding speed determines amount of energy input/unit length or, in other words, the time over which energy is applied at any particular point during ultrasonic welding [2]. Use of higher welding speeds minimizes sonotrode resident times and hence does not facilitate transfer of sufficient welding energy. Consequently, the magnitude of stresses generated at the interface would be insufficient to cause complete oxide layer removal and to induce adequate plastic deformation at surface contact points. This explains why the deposits showed decrease in LWD with increase in welding speed. Use of 28 mm/s, the lowest speed within the range selected in this study, was found to produce the best results. It is thus necessary to study the effect of welding speed at still lower levels in order to see if a further decrease in welding speed contributes to any further increase in LWD.

It is thus necessary to study the effect of welding speed at still lower levels in order to see if a further decrease in welding speed contributes to any further increase in LWD. To address this issue, experiments were conducted at four different speed levels below 28 mm/s, keeping all other parameters constant at their optimum levels. Observations from these follow-up experiments are discussed in Section 3.1.3.Effect of normal force: The LWD increased with increase in normal force from 1450 to 1750 N (Fig.7c). The applied normal force in ultrasonic welding not only serves to bring the mating surfaces in close contact with each other, but also determines the magnitude of dynamic interfacial stresses in combination with the oscillating shear forces due to ultrasonic vibration. An increase in applied normal force increases the magnitude of the resultant interfacial stresses701and hence, aids in bond formation.

An increase in applied normal force increases the magnitude of the resultant interfacial stresses701and hence, aids in bond formation. This explains why the LWD increased with increase in normal force from 1450 to 1750 N. Further increase in normal force to 1900 N, however, resulted in considerable drop in LWD (Fig.7c). While the exact reason for this behaviour is not clear at present, there are several potential explanations. Use of too high a normal force might result in excessive interfacial stresses leading to breakage of already formed bonds. Also, an increase in normal force will necessitate an increased sonotrode oscillatory force to maintain the same frequency.

LWD, (c) Normal force vs. LWD, (d) Substrate temperature vs. LWD.Effect of substrate temperature: LWD increased with increase in substrate temperature from 75 to 300(cid:113)F (Fig.7d). During ultrasonic welding, the in-situ raise in interface temperature as a result of friction plays a key role in bond formation by (i) reducing the flow stress of the material, (ii) enhancing atomic diffusion, and (iii) the driving force for recrystallization [3]. In addition, any strain hardening effect during plastic deformation would beincreasing70240300reduced at elevated temperatures.

In addition, any strain hardening effect during plastic deformation would beincreasing70240300reduced at elevated temperatures. Use of external thermal energy input in the form of elevated substrate temperature would further enhance these effects and thus promote bond formation during ultrasonic welding.3.1.3 Further experiments at lower welding speedsFollowing the initial round of experiments, additional experiments were conducted at four different speeds below 28 mm/s keeping all other parameters constant at their optimum levels. Table 5 shows the parameters used in these experiments along with the results of linear weld density measurements. Representative microstructures of the deposits made at 24 mm/s and 12 mm/s are shown in Fig.8. Microstructures of 20 mm/s and 16 mm/s deposits were quite similar to those of 24 mm/s and 12 mm/s deposits, respectively. All the four deposits showed only a few defects, mostly belonging to the point-like defect category.

All the four deposits showed only a few defects, mostly belonging to the point-like defect category. As can be seen, use of lower welding speeds resulted in further increase in %LWD, up to 98%, which is compared favorably to Run # 14, the best run in the initial round of experiments, with 90% LWD. Among the four deposits made in the follow-up experiments, samples made at 16 mm/s and 12 mm/s were found to be slightly better compared to the other two. Thus the results show that %LWD appears to asymptotically approach 100% as welding speed decreases.(a)(b)Fig.8. Microstructures of UC deposits (longitudinal section): (a) Welding speed: 24 mm/s (b) Welding speed: 12 mm/s. Other parameters were kept constant at their optimum levels (oscillation amplitude: 16 (cid:80)m, Normal force: 1750 N, and Temperature: 300(cid:113)F).

Other parameters were kept constant at their optimum levels (oscillation amplitude: 16 (cid:80)m, Normal force: 1750 N, and Temperature: 300(cid:113)F). Arrows indicate welding direction for each layer.3.1.4 Optimum process parametersBased on the observations made in the present study, the following was identified as the optimum parameter combination for alloy 3003: oscillation frequency – 16 (cid:80)m, welding speed – 16 mm/s, normal force – 1750 N, substrate temperature – 300(cid:113)F. A welding speed of 16 mm/s was chosen as a likely optimum condition, as further reduction in welding speed does not contribute to an increase in %LWD sufficient to justify the increase in fabrication time. Use of these parameters, barring other geometry-induced effects, should result in 98% linear weld density in Al alloy 3003 deposits.703Table 5. Welding speeds used in follow-up experiments and corresponding LWD results.

Welding speeds used in follow-up experiments and corresponding LWD results. Other parameters were kept constant at their optimum levels (oscillation amplitude: 16 (cid:80)m, Normal force: 1750 N, and Temperature: 300(cid:113)F).Welding speed (mm/s) 24 20 16 12LWD(%)94 96 98 98The experiments conducted in the present study do not facilitate statistical assessment of the optimum process parameters at room temperature. Operation at room temperature is important for the embedding of some temperature-sensitive components, such as electronics. When, in a future study, the process parameters are optimized for room temperature, it should be possible to achieve significantly higher weld density levels in the deposits than those observed in this study at room temperature, although not likely as high as those achievable at 300(cid:113)F.Another aspect that was not addressed in this study is interacting influences between two or more process parameters.

While the experimental matrix utilized in the present study does not facilitate precise identification of interactions, certain general trends may be noted. For example, use of low oscillation amplitude in combination with high welding speed and/or with low contact force may be seen to result in low weld density levels. Similarly, use of slower welding speeds and/or higher normal forces at a given oscillation amplitude may be seen to result in relatively higher weld density levels. Further work is necessary to precisely identify the interactions between various process parameters.Thus, while a high level of LWD can be achieved by appropriately choosing processparameters, this approach has certain limitations:(cid:120) While parameter optimization certainly helps minimize defect formation, defects cannot be eliminated altogether.

Thus, use of optimum parameters is not a complete solution.(cid:120) Low welding speeds significantly increase build time and overall cost of partfabrication.(cid:120) Elevated substrate temperatures put severe limitations on process capabilities. For example, parts with embedded electronics or other temperature-sensitive devices cannot be fabricated employing elevated substrate temperatures. In addition, alloys which age-harden at modest temperatures would see a change in bulk properties if processed at elevated temperatures.(cid:120) Use of high oscillation amplitude and/or normal force in combination with low welding speed can be damaging to the sonotrode [2,3]. This may necessitate frequent sonotrode cleaning or replacement.

This may necessitate frequent sonotrode cleaning or replacement. More importantly, the severe processing conditions can lead to unacceptable levels of strain hardening and/or fatigue-related effects at the interface, which could hamper bond strength and overall part mechanical properties, as shown in a previous investigation [4,5].Therefore, one cannot rely entirely on process parameters for ensuring a high %LWD in ultrasonically consolidated parts. It is thus necessary to look for alternative strategies which can result in 100% bonding even when processing at higher welding speeds and at ambient substrate temperature. Towards this end, a simple technique was devised, as described below.7043.2 Effect of surface machiningAs noted earlier, sonotrode-induced surface roughness was identified as the major source of defects in ultrasonically consolidated parts.

In order to minimize the incidence of defects due to surface roughness, a simple technique was devised incorporating an intermediate surface machining step after depositing each layer. Experiments were conducted at various welding speeds at both ambient and elevated substrate temperatures. Fig.9 shows the microstructures of deposits made employing the surface machining technique at various welding speeds at 300(cid:113)F. Similarly, Fig.10 shows the microstructures of deposits made employing the surface machining technique at various welding speeds at 75(cid:113)F. These microstructures conclusively demonstrate the effectiveness of surface machining for improving the LWD of ultrasonically consolidated parts.(a)(b)(c)Fig.9. Microstructures of Al alloy 3003 UC deposits with surface machining (longitudinal section).

Microstructures of Al alloy 3003 UC deposits with surface machining (longitudinal section). (a) Welding speed: 28 mm/s, (b) Welding speed: 36 mm/s, and (c) Welding speed: 40 mm/s (Oscillation amplitude:16 (cid:80)m, Normal force: 1750 N, and Temperature: 300(cid:113)F).705(a)(b)Fig.10. Microstructures of an Al alloy 3003 UC deposit produced with surface machining (longitudinal section) (a) Welding speed: 32 mm/s, (b) Welding speed: 40 mm/s (Oscillation amplitude:16 (cid:80)m, Normal force: 1750 N, and Temperature: 75(cid:113)F).One major benefit of surface machining is that high-quality deposits can be made at welding speeds that are significantly higher than those which results in high-quality deposits without surface machining. Welding speeds up to 36 mm/s at 300(cid:113)F and up to 32 mm/s at 75(cid:113)F showed nearly 100% LWD, with just a few isolated point-like defects.

Welding speeds up to 36 mm/s at 300(cid:113)F and up to 32 mm/s at 75(cid:113)F showed nearly 100% LWD, with just a few isolated point-like defects. Thus with surface machining, it is possible to achieve nearly 100% LWD even when processing at significantly higher welding speeds and even when processing at ambient temperature. However, defects can occur even with surface machining when the welding speed is too high, as illustrated in Fig.9c and Fig.10c, although LWD levels are considerably higher than those achievable with regular UC processing (without surface machining) under similar processing conditions.The observed improvement in bond formation utilizing the surface machining techniquecan be attributed to the following.i)ii)iii)iv)Removal of sonotrode-induced surface roughness facilitates intimate contact between mating surfaces, leading to a significant increase in number of surface contact points at which bonding occurs.

Surface machining removes the hills and valley pattern on the foil surface caused by the sonotrode motion, which would otherwise manifest into defects and entrapped air. Surface machining completely removes the oxide layer on one of the mating surfaces during ultrasonic welding, thus all problems associated with oxide layers are reduced to half. Surface machining removes any work hardened layer either completely or partially on the deposited foil surface. This would facilitate easier plastic flow at the interface when the subsequent layer is deposited, which is one of the important conditions for bond formation during ultrasonic welding.706The current work shows that the surface machining technique satisfactorily solves theproblem of defects in ultrasonically consolidated parts, with the following benefits:(cid:120) The technique widens the process window for satisfactory part fabrication.

(cid:120)It facilitates part fabrication at significantly higher welding speeds and/or at ambient temperatures without compromising on LWD.(cid:120) With surface machining, one can avoid the use of such process parameter selections that lead to excessive work hardening and/or fatigue-related effects at the interface.(cid:120) The current method is extremely simple and can be implemented without anymodification to the commercial UC equipment.(cid:120) The current work shows that, with proper choice of process parameters and surface machining, the ultrasonic consolidation process is quite capable of producing high quality parts on par with established wrought or cast processing techniques.It should be noted that the introduction of an intermediate surface machining step adds to the overall build time. In addition, if milling parameters are not optimized, an intermediate surface machining step can result in delamination of previously deposited layers.

In addition, if milling parameters are not optimized, an intermediate surface machining step can result in delamination of previously deposited layers. However, when appropriate milling parameters are used, the gains in part quality, welding speed and manufacturing flexibility may more than compensate for the extra machining time. Further, it is possible to significantly reduce intermediate machining times when a surface machining set-up is developed specifically suited to intermediate surface machining (as the current UC apparatus was not designed with this in mind).Alternatively, one may envision the use of other means of removing the surface roughness that could be more time-effective than machining, such as surface rolling, planishing, or acid or chemical etching. Finally, the current work raises important questions about ultrasonic sonotrode design strategies.

Finally, the current work raises important questions about ultrasonic sonotrode design strategies. Sonotrode designers, apart from other considerations, should pay greater attention to the problem of sonotrode-induced surface roughness. With improved sonotrode designs, this problem can be significantly reduced, possibly eliminating the need for surface machining altogether.4. ConclusionsBond formation during ultrasonic consolidation of Al alloy 3003 is strongly dependent on process parameters. Variations in oscillation amplitude, welding speed, normal force and substrate temperature contribute to statistically significant variations in linear weld density. Oscillation amplitude exerts the strongest effect. Linear weld density increases with increase in oscillation amplitude and normal force up to a certain level. Further increase in either oscillation amplitude or normal force results in a drop in linear weld density.

Further increase in either oscillation amplitude or normal force results in a drop in linear weld density. Linear weld density asymptotically approaches 100% with decrease in welding speed. Welding speed presents a clear trade-off between linear weld density and part build time. Use of elevated substrate temperatures promotes bond formation during ultrasonic consolidation. Linear weld density increases with increase in substrate temperature. The following combination of parameters produced the best results (98% linear weld density) during ultrasonic consolidation of Al alloy 3003 based on the parameter range studied: oscillation frequency – 16 (cid:80)m, welding speed – 16 mm/s, normal force – 1750 N, substrate temperature – 300(cid:113)F.707Sonotrode-induced surface roughness on the foil surface is a major source of defects in ultrasonically consolidated parts.

Li, Maximum height to width ratio of freestanding structures built using ultrasonic consolidation, Solid Freeform Fabrication Symposium, Austin, TX, 2006. G.D. Janaki Ram, Y. Yang and B.E. Stucker, Effect of process parameters on bond formation during ultrasonic consolidation of Al alloy 3003, submitted to Journal of Manufacturing Systems, April 2006.708.

proving Tensile Mechanical Properties of FDM-Manufactured Specimens via Modifying Build ParametersMohammad Shojib Hossain, Jorge Ramos, David Espalin, Mireya Perez, Ryan Wicker W.M. Keck Center for 3D Innovation, Department of Mechanical Engineering, The University of Texas at El Paso, 500 W University Ave, El Paso, TX79968AbstractIn this paper, the focus was on improving tensile mechanical properties of FDM- manufactured parts by adjusting FDM processing parameters and analyzing stress concentration features between adjacent roads of material. FDM processing parameters are specified by the user via Insight – the file preparation software for most FDM machines. Even though Insight gives the impression that adjacent roads are to be deposited and connected throughout, an optical imaging observation of the deposited material revealed that adjacent roads are not consistently connected forming voids that reduce mechanical performance.

Therefore, this work reports the tensile mechanical properties of specimens built using three sets of parameters: standard/default parameters, an Insight revision method, and a visual feedback method. When compared to the default build parameters, the experimentally determined, visual feedback approach produced specimens, in some cases, exhibiting as high as 19% improvement in ultimate tensile strength.IntroductionThe material extrusion additive manufacturing (AM) process [1], commonly known as fused deposition modeling (FDM) is well known for its use in producing prototypes for concept modeling, pre-surgical models in medical applications, and various other uses. Now, efforts have been focused on transforming the technology towards manufacturing production-grade and end- use products [2].

Now, efforts have been focused on transforming the technology towards manufacturing production-grade and end- use products [2]. With the current expansion of different material extrusion AM machines, such as the FDM series including the Fortus models and uPrint (Stratasys Inc., Eden Prairie, MN), Makerbot ReplicatorTM (MakerbotTM, Brooklyn, NY ), and numerous other desktop machines, there are various options for processing parameters such as build orientation, raster angle, contour width, raster width, raster to raster air gap, and more. Some software packages have limited access to the processing parameters, by only providing access to build styles (e.g., sparse, solid), hindering the effort to perform process optimization due to the lack of access to raster widths and contour widths, for example. Along with other manufacturing processes, FDM process parameters play a significant role in the performance of the fabricated parts, especially in regards to mechanical properties.

Along with other manufacturing processes, FDM process parameters play a significant role in the performance of the fabricated parts, especially in regards to mechanical properties. For a better understanding of these parameters, a brief description of the FDM operation is provided here.The main principle of FDM technology is to generate parts directly from three dimensional computer aided design (CAD) data by using a material extrusion process. In most cases, final parts do not require machining. First, a three dimensional CAD model is created and sent to the specific AM software package. The software package used for FDM and Fortus machines is Insight (Stratasys Inc., Eden Prairie, MN). The model is sliced layer by layer withinthe Insight job file. Then, toolpaths are generated based on user-specified parameters, which are described here. A rendering of the toolpaths is given by the software.

Then, toolpaths are generated based on user-specified parameters, which are described here. A rendering of the toolpaths is given by the software. After tooling information, created by Insight, is sent to the FDM machine, a thermoplastic filament is driven into the heated liquefier (Fig. 1) where the plastic reaches a flowable state, and is extruded through a small diameter tip. During deposition of material, the tip moves in the X and Y directions, to create a layer consisting of contour and rasters. Contour creates the peripheral shape of the 3D model and raster is the internal fill pattern in-between the contours. Normally, to hold the whole part in place on the platform, and for building overhanging features, a sacrificial support material is used, which can be broken away or dissolved in a cleaning solution, once the part is completely built.

After creating one layer, the bed drops down a distance equal to the height of the layer and a new layer is then deposited over the previous one. This process of building layers subsequently goes on, until the whole part is complete. To maintain a constant heated environment, the entire process is done in a closed chamber which is also known as the envelope. This heated environment helps to improve the interlayer bonding [3] and reduces shrinking, warping and internal stresses [4].The FDM processing parameters include build orientation, raster angle (RA), contour width (CW), number of contours, raster width (RW), raster to contour air gap, raster to raster air gap (RRAG) and slice height. In this paper, RA, CW, RW, and RRAG were varied using one specific build orientation. The build orientation is the orientation of the part with respect to build platform.

In this paper, RA, CW, RW, and RRAG were varied using one specific build orientation. The build orientation is the orientation of the part with respect to build platform. According to the ASTM F2921-11 standard, the coordinate system of the build platform and the specific build orientation, in this case XYZ, are shown in Fig. 2. RA is the angle created by the raster and the positive X direction of the build platform. CW and RW is the width of contour and raster, respectively. RRAG is the distance between the edges of two adjacent rasters (Fig.3). A negative RRAG (obtained by decreasing RRAG from zero) causes the partial overlap between two adjacent rasters.Fig. 1 Schematic diagram of FDM machineFig. 2 XYZ build orientation used for the fabrication of ASTM D638 Type I specimens. The XYZ terminology is based on ASTM F2921-11Fig.

1 Schematic diagram of FDM machineFig. 2 XYZ build orientation used for the fabrication of ASTM D638 Type I specimens. The XYZ terminology is based on ASTM F2921-11Fig. 3 FDM Build ParametersThe choices of commercially available thermoplastic materials, for the use with FDM includes acrylonitrile butadiene styrene (ABS), polycarbonate (PC), a PC-ABS blend, polyphenylsulfone (PPSF), ULTEM, and several varieties of the aforementioned materials designed for biocompatible or static dissipation applications [5]. Prior literature also mentioned, other materials processed by FDM including polycaprolactone (PCL) [6], polymethyl methacrylate (PMMA) [7], and composites that contain ceramic, glass, or metal fillers [8, 9]. In this experiment, PC was chosen because of its common use with FDM. Benefits of the material include its high ultimate tensile strength (UTS) (approximately twice of ABS), high hardness, and high toughness.

Benefits of the material include its high ultimate tensile strength (UTS) (approximately twice of ABS), high hardness, and high toughness. Prior works showed similar testing mostly with ABS, however a data set for PC can be of great use to the FDM community.Prior efforts have been made towards the modification of FDM-build parameters for improvement in mechanical properties. A Study by Masood et al. showed that the UTS of PC decreased about 8% and 13% using two different RWs of 0.4064 mm and 0.6064 mm respectively, while changing the RA from 45º/-45º to 90º/0º [10]. In the same study, the increase of UTS was around 18% by changing the RA from 45º/-45 º to 90º/0 º using a RW of 0.8314 mm. The highest UTS obtained in that experiment was 58.8 MPa using a RA of 45 º/-45 º and a RW of 0.6064 mm. Ahn et al.

The highest UTS obtained in that experiment was 58.8 MPa using a RA of 45 º/-45 º and a RW of 0.6064 mm. Ahn et al. tested the same idea of changing the RA, using ABS as the build material and found an increase of about 8%, by changing the RA from 45º/-45 º to 90º/0º using RWs of 0.508 mm and 1 mm [11]. In their tests, there was less than a 2% change in UTS using different RWs. In addition, they detected increased build times using small RWs. On the other hand, using ABS, Sood et al. observed that, a change in RA from 60º/-30º to 90º/0º led to a decrease in UTS by approximately 19% [12]. The author stated, this was because a 0º RA yielded the longest rasters which led to an increase in stress accumulation along the deposition direction of ABS, resulting in higher distortion and weak bonding. Regarding RRAG on UTS, Ahn et al. revealed that a RRAG of -76.2 µm (-0.003 in) increased the tensile strength of ABS by around 30% using a RA of 0º/90º in comparison to a RRAG of 0 mm [11].

Regarding RRAG on UTS, Ahn et al. revealed that a RRAG of -76.2 µm (-0.003 in) increased the tensile strength of ABS by around 30% using a RA of 0º/90º in comparison to a RRAG of 0 mm [11]. Negative air gaps developed dense structures which led to greater UTS. On the contrary, the study by Sood et al. revealed that a positive RRAG caused the material to flow between adjacent layers, increasing the bonding between surfaces [12], which led to an increase of around 15% when building with thinner layers, such as 0.127 mm. Changing layer thickness to 0.254 mm, showed no significant difference in UTS using different RRAGs. However, it was unclear if the flow between layers was the result of other factors, such as heat dissipation.As stated above, a survey of the literature revealed prior work that has been done to improve the mechanical properties through the modification of build parameters.

The work presented here emphasizes on establishing a method, which includes optical feedback, to expose the reason for improvement in mechanical properties of PC specimens: UTS, Young’s modulus, and tensile strain. Throughout the experiment, it was revealed that the removal of air gaps between adjacent rasters and between contours and rasters, appeared to have a positive effect on mechanical properties. Two methods for modifying the build parameters were developed by considering the toolpath renderings and by examining low-magnified optical images of the test specimens. These methods were compared with default build parameters. It was revealed that the UTS could be improved as high as 19% compared to specimens built using default parameters.

These methods were compared with default build parameters. It was revealed that the UTS could be improved as high as 19% compared to specimens built using default parameters. Overall, a maximum UTS of 53.75 MPa was attained, as compared to 46.84 MPa using default parameters.Experimental ProcessParameter ModificationsThe test specimen was created according to the ASTM D638 Type I specimen [13] using SolidWorks (Dassault Systemes SolidWorks Corp, Velizy, France). The XYZ build orientation was selected for all the test specimens. For this orientation, three different RAs were used: 0º/90º, 30º/-60º, and 45º/-45º. Three methods were tested to establish the build parameters:Default methodInsight revision method Visual feedback methodFor the default method, the building parameters were obtained from the default value provided by Insight. Gaps shown in optical images revealed the need for parameter modification.

Gaps shown in optical images revealed the need for parameter modification. For the Insight revision method, the job files created via Insight software were analyzed. The CW and RW were changed from 0.4064 mm to 0.8128 mm in increments of 0.0254 mm using a 0º/90º RA. From all the combinations of CW and RW, the rendered image showing the smallest gaps between adjacent rasters and contour was selected for building. After that, those parameters were applied to the other two RAs (30º/-60º and 45º/-45º). The gaps between rasters and contour were much more apparent in the rendering for a 0º/90º RA. Therefore, this particular RA was chosen as a modification basis. The parameters that showed better renderings for the 0º/90º RA had a tendency to yield good results for the other RAs. For the visual feedback method, the specimens built with the Insight revision method were analyzed with a microscope.

For the visual feedback method, the specimens built with the Insight revision method were analyzed with a microscope. Gaps were visible between two adjacent rasters, though these were not identifiable in the rendering created by Insight. Five different RRAGs, from -0.0025 mm to -0.0254 mm, were analyzed with the microscope. As optical images for a -0.013 mm RRAG displayed an absence of air gaps between two adjacent rasters, it was selected as the best value. The building values are shown in Table 1.Specimen Construction:The specimens were built with a Fortus 900 mc (Stratasys Inc., Eden Prairie, MN). According to the manufacturer’s specifications, this machine has a build chamber of 914 mm × 610 mm × 914 mm. PC support was used as support material. The model material, support material, and envelope temperatures were 345ºC, 365ºC, and 140ºC respectively (these are standard model, support, and envelope temperatures for the PC and PC Support materials provided by Stratasys).

The layer thickness or slice height was set at 0.254 mm.Optical evaluation:A Retiga 2000R fast charge-coupled device camera (QImaging, Surrey, Canada) installed on a Leica MZ16 stereomicroscope (Leica Microsystems, Wetzlar, Germany) was used to take low magnification optical images (32x). The optical images were analyzed with the aid of QCapture Pro software (QImaging, Surrey, Canada).Table 1: Building ParametersBuild orientation XYZRaster angle 0ᵒ/90ᵒ 30ᵒ/-60ᵒ 45ᵒ/45ᵒDefault Parameter RW 0.508 0.508 0.508RRAG 0 0 0 *CW= Contour Width, RW= Raster Width, RRAG= Raster to Raster Air gapCW 0.508 0.508 0.508Insight revision method RRAG CW RW 0 0.432 0.432 0 0.432 0.432 0 0.432 0.432Visual feedback method RRAG RW CW -0.013 0.432 0.432 -0.013 0.432 0.432 -0.013 0.432 0.432** All the values are in mmMechanical Testing:Tensile tests were carried out by using an Instron 5866 (Instron, Norwood, MA) tensile testing machine, which had a load cell of 10 kN.

The load cell measurement accuracy was ± 0.4% (as per manufacturer’s specifications). The ramp speed was kept to a constant 5 mm/min. The UTS, Young’s modulus, and tensile strain were calculated by the built-in software, Bluehill (Instron, Norwood, MA). Five specimens (totaling 45) were tested for every parameter modification. Before mechanical testing, all specimens were conditioned for 40 hours at 23 ºC and 50% relative humidity, according to ASTM standard D618 [14].Results and discussionOptical Observations:Fig. 4 shows the optical images of rendered toolpaths created by Insight and actual material deposition. In the Fig. 4, using the default parameters, the gaps between contour and rasters are visible (Fig. 4a and 4b). Fig. 4c and 4d show that, using modified parameters (Insight revision method, contour width at 0.432 mm and raster width at 0.432 mm), gaps between the contour and rasters were removed.gap between raster and contour(a)(b)no gap(c)(d)Fig.

4 Rendered toolpaths generated by Insight (left) and actual deposited material (right) for the XYZ orientation using a 0ᵒ /90ᵒ raster angle; a) and b) default parameters (CW 0.508 mm, RW 0.508 mm, RRAG 0 mm), c) and d) Insight revision method parameters (CW 0.432 mm, RW 0.432 mm, RRAG 0 mm)Despite the efforts to remove air gaps with Insight revision, some gaps could be seen which were not visible in Insight renderings. Fig. 5 shows more detailed observations of the three methods used: default method, Insight revision method, and visual feedback method. This figure shows the optical imaging observations for the XYZ build orientation using RAs at 0º/90 º, 30º/-60º, and 45º/-45º. Gaps can be seen between adjacent rasters using default values (CW at 0.508 mm, RW at 0.508 mm, and RRAG at 0 mm). Using Insight revision, CW and RW were decreased, (CW at 0.432 mm, RW at 0.432 mm, and RRAG at 0 mm) and gaps between rasters were reduced.

Using Insight revision, CW and RW were decreased, (CW at 0.432 mm, RW at 0.432 mm, and RRAG at 0 mm) and gaps between rasters were reduced. The gaps were completely removed using visual feedback modifications (CW at 0.432 mm, RW at 0.432, and RRAG at -0.013 mm).Default (CW:0.508 RW:0.508 RRAG:0)Insight revision method (CW:0.432 RW:0.432 RRAG:0)Visual feedback method(CW:0.432 RW:0.432 RRAG: -0.013)RA 0º/90º500 um500 um500 um.RA 30º/-60º500 um500 um500 umRA 45º/-45º500 um500 um500 umFig. 5 Optical images of specimens built in the XYZ orientation with airs gaps highlighted by black circles (parameters are set in mm)Mechanical Testing ResultsThe UTS results obtained for the three RAs are presented in Fig. 6. For all the RAs, the greatest increase in UTS was obtained with the visual feedback method, compared to the UTS obtained using default parameters.

6. For all the RAs, the greatest increase in UTS was obtained with the visual feedback method, compared to the UTS obtained using default parameters. The negative RRAG led to a decrease in gaps between rasters, which is the reason for the increase in UTS. This is supported by Ahn et al. who indicated negative RRAGs led to stronger, denser, structures [11]. The highest percent increase in UTS compared to the default was 19% (44.76 MPa vs. 53.22 MPa), obtained using a 30ᵒ/-60ᵒ RA. The percent increases for all three RAs are listed in Table 2.A two sample t-test showed that UTS results acquired with the default parameters and visual feedback method were significantly different (p value < 0.05). The t-test results are listed in Table 3. The ANOVA test (table 4) shows that UTSs, among different RAs, are not significantly different (p value > 0.05) using same build method.

The t-test results are listed in Table 3. The ANOVA test (table 4) shows that UTSs, among different RAs, are not significantly different (p value > 0.05) using same build method. Whenever the RA was changed, the delta angle (change of raster angles between two adjacent layers) was kept constant at 90º. Thus, the directional effect of one layer might be nullified by the delta angle of the adjacent layer. This might be the reason behind the lack of significant differences in UTS when changing RA. However, this statement is not supported by Masood et al. [10] who observed a change in UTS using different RAs. Their observations were not statistically analyzed.Fig. 6 UTS for XYZ build orientation using three different RAs.

7 shows the Young’s modulus for XYZ build orientation. The highest average Young’s modulus achieved was 1816 MPa. This was obtained using a 0º/90 º RA with the visual feedback method (compared to 1595 MPa with the default). Analysis of the optical images, led to the understanding that changes in Young’s modulus may be directly related to removal of air gaps. For example, in Fig. 5, for all three RAs, removal of gaps is visible with the visual feedback method, compared to other two methods. Hence, there is an increase in Young’s modulus.Fig. 7 Young’s modulus for XYZ build orientation. Each bar represents the average of at least 5 specimens and the error bars are +/- standard deviationFig. 8 represents the tensile strain values for the XYZ build orientation using the three RAs. The highest average tensile strain obtained was just over 8% for the 45º/-45º RA with the Insight revision method. The values typically varied from 3.5% - 8%. Nadooshan et al.

The highest average tensile strain obtained was just over 8% for the 45º/-45º RA with the Insight revision method. The values typically varied from 3.5% - 8%. Nadooshan et al. reported a maximum tensile strain of 5% using PC [15]. With the modification in build parameters, higher values of tensile stain are achievable.Fig. 9 shows the characteristic stress-strain curves for the above mentioned results. It illustrates the curves closest to the average values obtained for the 0º/90º RA using the default, Insight revision, and visual feedback methods. The figure only illustrates the 0º/90º RA as there was no significant difference in UTS amongst different RAs. This graph helps to get an understanding of the benefits achievable from performing the methods discussed in this paper.Fig. 8 Tensile strain for XYZ build orientation. Each bar represents the average of at least 5 specimens and the error bars are +/- standard deviationFig.

8 Tensile strain for XYZ build orientation. Each bar represents the average of at least 5 specimens and the error bars are +/- standard deviationFig. 9 Characteristic stress-strain curves for three methods using 0º/90º RA. Default method is represented by a solid line (blue), Insight revision method by a round dotted line (green), and visual feedback method by a dashed line (purple).ConclusionBuilding parameters (including RA, RW, CW, and RRAG) play a vital role in improving mechanical properties of FDM-produced parts. In all circumstances, parameter modification using the Insight revision method, improved UTS compared to default values. The visual feedback method further improved UTS, introducing negative RRAGs, which led to an average increase in UTS of 16 % compared to the specimens built with default parameters. Overall, the highest average value of UTS obtained for PC was 53.75 Mpa (compared to 46.84 MPa with the default).

Overall, the highest average value of UTS obtained for PC was 53.75 Mpa (compared to 46.84 MPa with the default). Optical imaging, or the visual feedback method, provided great assistance in improving mechanical properties. Optical observations also led to the realization that what was rendered by the insight job file, was not necessarily what was produced. Air gaps were present in the final specimens, despite that those gaps were not visible in the job file created by Insight. Therefore, a separate calibration protocol or an in-situ monitoring system can identify necessary adjustments to the processing parameters. Introducing negative RRAGs created stronger inter-raster bonding and minimized or even removed gaps (at least through optical imaging observations) between rasters. Future work may include the effect of build orientation as well as layer thickness in further improving mechanical properties of FDM parts. However, modifications in layer thickness may compromise building time.

However, modifications in layer thickness may compromise building time. While the visual feedback method does reduce the gaps in a part, it does not completely eliminate them. Therefore, measurements can be conducted in the future to determine how the changes in porosity of the samples, through the elimination of air gaps, affect density.AcknowledgementThe research presented here was conducted at The University of Texas at El Paso (UTEP) within the W.M. Keck Center for 3D Innovation (Keck Center), recently expanded to over 13,000 sq. ft. The W.M. Keck Center provides access to state-of-the-art facilities and equipment as a result of funding from the State of Texas Emerging Technology Fund. Financial support for the research was provided, in part, by the University of Texas System Louis Stokes Alliance for Minority Participation Program under grant NSF-HRD-1139929 as well as the Mr. and Mrs. MacIntosh Murchison Chair I in Engineering Endowment.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperImproving the Mechanical Response of the IWP Exo-skeletal Lattice Through Shape OptimizationJoseph W. Fisher1,3, Simon W. Miller3, Joseph Bartolai3, Timothy W. Simpson1,21Department of Mechanical Engineering, The Pennsylvania State University, University Park, PA 16802 2Department of Industrial and Manufacturing Engineering, The Pennsylvania State University, University Park, PA 16082 3Applied Research Lab, The Pennsylvania State University, State College, PA 16803AbstractTriply Periodic Minimal Surfaces have been identified as good candidates for the generation of lattice structures produced with additive manufacturing. These TPMS-based lattice structures avoid sharp features that are characteristic of strut-based lattice structures because of their constant zero mean curvature.

These TPMS-based lattice structures avoid sharp features that are characteristic of strut-based lattice structures because of their constant zero mean curvature. Although studies have explored part-scale optimization using TPMS-based lattice structures, they have only varied the volume fraction by changing the level set in the approximate surface equations. By defining new parameterizations in the approxi- mate surface equation, we can redistribute volume within the lattice structure at any volume fraction. In this paper, we introduce an approach for optimization of this new parameterization of TPMS equations using the Borg multi-objective evolutionary algorithm. We demonstrate this framework on the IWP exo-skeletal lattice under uniaxial compression. A relationship between the new parameters and the level set is derived for designs on the Pareto frontier of the optimized IWP TPxS.

A relationship between the new parameters and the level set is derived for designs on the Pareto frontier of the optimized IWP TPxS. The performance of the Pareto optimal designs and the efficacy of the optimization approach are shown by comparing to the standard IWP lattice and four other lattices that share the same topology. The optimized designs are implemented and shared in custom nTopology blocks.Keywords: Lattice Structures, Triply Periodic Minimal Surface, Shape Optimization1INTRODUCTIONOne of the major draws of additive manufacturing (AM) is that components can be redesigned to reduce weight with the same or improved performance by taking advantage of the design freedom offered by AM [1]. One method of reducing weight in parts designed for AM is by incorporating lattice structures into designs. Lattice structures, also known as lattices, are repeating three dimen- sional structures that can be used to fill regions of the component.

Lattice structures, also known as lattices, are repeating three dimen- sional structures that can be used to fill regions of the component. These structures exist between the macro-scale structure of the component and the microstructure of the material, unlocking new design opportunities. A lattice offers the ability to generate regions that are not completely solid but retain a significant portion of the mechanical properties of the material it is made from, thus reducing the resulting part’s weight [2, 3]. Many studies have proposed a large variety of different1953(a)(b)(c)(d)Figure 1: Models of (a) the Body Centered Cubic (BCC) unit cell and (b) the I-graph–Wrapped Package graph (IWP) exo-skeletal (TPxS) lattices which share the same topology. Finite element analysis (FEA) of each lattice in uniaxial compression is presented in (c-d) with a 13% reduction in peak von Mises stress in the IWP TPxS compared to the BCC.

Finite element analysis (FEA) of each lattice in uniaxial compression is presented in (c-d) with a 13% reduction in peak von Mises stress in the IWP TPxS compared to the BCC. The stress concentrations where struts in the BCC lattice intersect is avoided by the smooth curvature of the IWP TPxS. The boundary conditions and loads for the FEA are identical to those of the validation models which are given in Table 1.lattice topologies [e.g., 4, 5, 6]. Two classes of lattice structures have recently seen significant traction within the field, those based on struts and those based on Triply Periodic Minimal Surfaces (TPMS) [7]. A strut-based lattice is a three dimensional truss like structure composed of individual struts that join together at discrete nodes. One limitation of strut-based lattice structures is that the sharp intersections between struts result in stress concentrations.

One limitation of strut-based lattice structures is that the sharp intersections between struts result in stress concentrations. These concentrations can be minimized by applying rounds to the junctions, but this can be computationally expensive when the lattice structure has many members [8, 9]. TPMS-based lattices are modeled using approximate equations defined by 3D Fourier series fits of surfaces of zero mean curvature (minimal) that exhibit periodicity in R3 (triply periodic) [10]. By definition, a minimal surface cannot have sharp corners, as it would violate the zero mean curvature. This property reduces stress concentrations, allowing for better utilization of material. The majority of strut and TPMS lattice structures vary greatly from one another, however, the strut-based Body Centered Cubic (BCC) and TPMS-based I-graph– Wrapped Package-graph (IWP) exo-skeletal (TPxS) lattice structures share the same connectivity (See Figures 1a and 1b).

As shown in Figures 1c and 1d, the IWP TPxS shows improvements to the mechanical properties compared to the BCC, which is consistent with Zhao et al. [11]. In addition, using a TPMS as a basis can be an efficient way to generate large and variable lattice structures [12].The Fourier series fit used to model lattices based on the IWP TPMS, given in Equation 1, was derived by Von Schnering and Nesper [10]. In this work we use the following shorthand for the where i is a spatial coordinate (i.e. x, y, z), di is the period of the cosine function Cni = cos lattice’s unit cell in direction i, and n is the number of periods of the trigonometric function in one period of the lattice in direction i. (e.g.

(e.g. 1, 2), when n is one, it is omitted simplicity.() i2πn di2 (CxCy +CyCz +CzCx) − 1( C2x +C2y +C2z)= t1954(1)TPMSEndo-skeleton (TPnS) Surface (TPSf) Exo-skeleton (TPxS)CombinedFigure 2: The IWP surface and the three types lattice structures modeled from it using Equation 1. Reproduced with permission from [13].The two terms on the left hand side of Equation 1 define the surface that approximates the IWP TPMS, and t is a level set that provides an offset to that surface allowing the volume of material in the lattice to be changed.Implicit equations that approximate TPMS, including Equation 1, can be used to model three unique types of lattice structures by defining solid material based on the value of the field in space. The three types are: the endo-skeleton (TPnS), the exo-skeleton (TPxS), and the surface (TPSf) [13].

The three types are: the endo-skeleton (TPnS), the exo-skeleton (TPxS), and the surface (TPSf) [13]. Respectively, these three lattice structures are modeled by defining the following regions of the field as solid: less than t, greater than t, and those between t and −t. We show these three lattices for the IWP in Figure 2, highlighting the unique geometries. For this work, we focus on the TPxS.For the discussion of lattice structures, we define the volume fraction as the volume of the solid material in the lattice normalized by the volume of the space that the lattice is filling. Lattice structures are often discussed in terms of their volume fraction in the literature as it provides a common basis for comparison between lattices that is independent of lattice topology and the geometry of the space that the lattice is filling.The first two terms of Equation 1 can be inspected separately to understand the impact that each has on the topology of the resulting surface.

Structures produced using each term alone are shown in Figure 3. Informally, we can say that the two terms respectively draw mass toward and away from the “nodes” of the lattice at the center and corner of the unit cell. If the second term in Equation 1 had a coefficient greater than one, the “nodes” would be smaller and the “beams” thicker. In Equation 1, the ratio of the two is fixed as 2:1, which provides a best fit to the IWP TPMS when t = 0 [10]. Since this equation was derived at a single level set (which is also a single volume fraction), and not based on any criteria specific to lattice design, it is likely that variation on Equation 1 can improve the performance of the lattice. With this motivation, we define a new equation with variables in place of the coefficients 2 and 1, shown below in Equation 2.

With this motivation, we define a new equation with variables in place of the coefficients 2 and 1, shown below in Equation 2. We also define our first research question: Can the response of the IWP TPxS lattice be improved at one or more volume fractions by varying the coefficients A and B present in its equation?A (CxCy +CyCz +CzCx) − B( C2x +C2y +C2z)− t > 01955(2)(a)(b)Figure 3: Models generated with (a) the first and (b) second terms of Equation 1 in isolation. The balancing of the two terms provides a useful approximation of the IWP TPMS. Both terms show mirror symmetry in x, y, and z through the center of the unit cell.The values of A, B, and t can be varied to generate unique structures, producing a 3D design space in which to search for optimal designs.

The optimization of TPMS-based lattice structures has been performed at the component scale by changing value of the level set t to redistribute mass within the lattice to meet the performance requirements of the component [14, 15, 16]. In addition, the optimization of strut-based unit cells has appeared multiple times in the literature with differ- ent approaches. Ground structure approaches, utilizing a multi-objective evolutionary algorithm (MOEA), have been demonstrated for generating novel unit cells for specific applications [17]. Several optimizations have been performed on the BCC lattice to improve the response of the lattice. Wang et al. [8] approached the optimization by adding rounds to the intersection of the struts whose radius was a function of the strut radius. Lee, Zhang, and Gu [18] applied a machine learning approach to optimize the profile of the struts which they parameterized as 30th order Bézier curves. Zhao et al.

Lee, Zhang, and Gu [18] applied a machine learning approach to optimize the profile of the struts which they parameterized as 30th order Bézier curves. Zhao et al. [19] defined the BCC lattice using struts that have a hexagonal cross-section that taper linearly from the nodes to the center of the beam. This approach reduces the complexity of the model, replacing curved surfaces with trapezoidal faces that are more easily meshed; however, this approach does result in different stress concentrations at the mid-span of the struts where the planer faces intersect. We show examples of the geometry generated using these three approaches in Figure 4. When discussing these different lattices in this work, we will reference them by the last name of the first author (i.e. Wang, Lee, and Zhao). These three optimizations of the BCC lattice all included uniaxial compression as a load case in their optimization, giving a common point of comparison.

Wang, Lee, and Zhao). These three optimizations of the BCC lattice all included uniaxial compression as a load case in their optimization, giving a common point of comparison. This leads to a second research question: Do the results of optimization of the IWP TPxS outperform optimizations of the BCC lattice which share the same topology? For TPMS-based lattices, no studies have performed this type of optimization on the unit cell prior to utilizing it at the component scale. As demonstrated by prior works (e.g., [8]), such an optimization1956(b) Figure 4: Optimized BCC unit cells from (a) Wang et al. [8], (b) Lee, Zhang, and Gu [18], and (c) Zhao et al. [19] at volume fractions of 0.1970, 0.2409, and 0.1977, respectively. The smooth curvature of (b) shows a strong similarity to the IWP TPxS; however, creases at the intersections of the struts persist through the optimization that lead to stress concentrations.(a)(c)can increase the performance of the lattice structure.

In the new parameterization in Equation 2, the coefficients A, B, and the parameter t can be used as inputs for the optimization of the lattice since they are not fixed values. Unit cells that have been optimized can then be utilized in component level optimization. In addition to other optimizations including uniaxial loading cases, the majority of lattice structure analysis and mechanical testing is performed in uniaxial compression [20, 4, 16]. For consistency, we target the same loading conditions. By applying periodic boundary conditions, we can approximate the response of an infinite lattice using a single unit cell [21].The BCC optimizations mentioned previously resulted in geometry that is defined either for any volume fraction or for only discrete volume fractions. Specifically, Wang et al. [8] and Zhao et al.

Specifically, Wang et al. [8] and Zhao et al. [19] developed mathematical relationships which can be used to model lattices at any volume fraction, where Lee, Zhang, and Gu [18] developed lattices only at discrete volume fractions. The former of these two is advantageous because they can be used to smoothly transition between different volume fractions when performing part scale optimization as demonstrated by [22, 8, 15].

Comparing these types of results motivates an additional research question: Do relationships between the optimal coefficients at different volume fractions exist, leading to a reduction of parameters?For each research question, we hypothesize the results below.RQ1: Can the response of the IWP TPxS lattice be improved at one or more volume fractions by varying the coefficients A and B present in its equation? We hypothesize that lattice performance can be improved, and at all volume fractions, by changing the weighting of A and B with an appropriate selection of t.RQ1.1: Do relationships between the optimal coefficients at different volume fractions exist, leading to a reduction of parameters? We hypothesize that such relationships exist1957between coefficients of true Pareto-optimal solutions, but the recovery of such rela- tionships is contingent on the objective evaluation approach and optimality criteria.

We note that optimization can only approximate the Pareto frontier [23, p. 17], so any relationships we find between coefficients will depend on this approximation.RQ2: Do the results of optimization of the IWP TPxS outperform optimizations of the BCC lattice which share the same topology? As we have shown (for a single volume fraction) the IWP TPxS out performs the BCC lattice in uniaxial compression (see Figure 1) and we hypothesize that optimization of the IWP TPxS will outperform optimization of the BCC provided both maintain the same topology.To answer these questions, we have developed a method of optimizing the unit cells of the IWP TPxS lattice. In Section 2, we present the definition of the design space and the problem definition; we then define the objective function and the method we use to evaluate it before discussing In Section 3, we evaluate relationships between the the selection of our optimization scheme.

Pareto-optimal designs and compare the Pareto-optimal lattices to existing lattices and discuss our research questions. We then apply homogenization approaches to study the lattice performance and anisotropy before summarizing our work and discussing possible extensions in Section 4.2 METHODSIn this section, we outline our approach to defining out optimization problem, the optimization method employed, and the implementation of the methods.2.1 Design Space & Problem DefinitionThe design space for our optimization problem is a real three dimensional space consisting of the parameters A, B, and t from Equation 2. We bound this space to a right rectangular prism by defining the following inequality constrains:0 ≤ A ≤ 1 0 ≤ B ≤ 1 −1 ≤ t ≤ 1Negative values of A or B were found to produce topologically distinct structures that do not share connectivity with the IWP TPxS. Therefore, we consider only positive values.

Therefore, we consider only positive values. Not all designs within the space defined by Equations (3) to (5) are valid. To fully define the feasible region of the design space, we first evaluate Equation 2 over a linearly-spaced 3D grid of 128,000 designs (a 40 × 40 × 80 grid). We evaluated each point for two criteria: connectivity and topology. To test connectivity, each design Equation 2 is evaluated over a 3D grid in x, y, z space. Taking a level set of this grid at t, we apply Lewiner’s Marching Cubes algorithm to define the surface of the lattice geometry as a triangular mesh [24]. The faces of this mesh are then used to construct an undirected graph. If all vertices of this graph are connected, the geometry of the design is connected also. We1958(3)(4)(5)(a)(b)(c)(d)Figure 5: Example structures that are possible outside the feasible region.

We1958(3)(4)(5)(a)(b)(c)(d)Figure 5: Example structures that are possible outside the feasible region. (a) a design that is discontinuous, and (b-d) connected designs with whose topology is inconstant with the IWP TPxS, because the second term of Equation 2 dominates to an increasing degree. The structure in (b) has internal voids at both the center and corners of the unit cell.note that this is not universally true for all unit cells of all TPMS-based lattices, such as the common unit cell of the gyroid which appears to have disconnected volume “floating” in the corners [13]. However, the “floating” volumes in the gyroid connects to the rest of the structure through adjacent unit cells. For the lattices of the IWP surface, the volume remains connected inside the unit cell, so the connectivity check described previously can be applied. To test the topology of the lattice, we evaluated Equation 2 at a corner point of the unit cell.

To test the topology of the lattice, we evaluated Equation 2 at a corner point of the unit cell. To match the desired topology this point must fall within the lattice. Examples of four designs that fall outside the feasible region are shown in Figure 5.For valid designs, we evaluated two metrics of interest: volume fraction and minimum hy- draulic diameter of the “beams” using Equation 6. Where Ac and p are the cross-sectional area and perimeter of the “beam” at the mid-span of the “beam”. We evaluated Ac and p using the surface mesh generated using the Marching Cubes algorithm. The hydraulic diameter is used in Section 2.1.2 to highlight the variation between the geometry of different valid designs independent of volume fraction.dh = 4AC pThe design space was visualized using the ARL Trade Space Visualizer1[25] to plot each valid design as a cube colored by the volume fraction of the design (see Figure 6).

The design space is an irregular octahedron, five of whose faces are already defined by Equations (3) to (5). Constraints for the remaining three faces were defined using the equation of a plane in 3D, commonly expressed as:ax + by + cz − d = 0where x, y, and z are the design variables A, B, and t, respectively. We compute the coefficients a, b, c, and d for each of the three surface using a modified version of the MATLAB function1http://www.atsv.psu.edu/whatisatsv.html1959(6)(7)Figure 6: Exploration of the 3D design space resulted in a polyhedron with 8 faces representing the feasible region (Ω). Color indicates the volume fraction which is the percentage of the lattice unit cell that is solid.plot_plane [26] and three noncollinear design points on each face of the domain. The coefficients a, b, and c respectively are the A, B, and t components of the plane’s unit normal, and d is the offset from the origin in the direction of the normal.

[27] explored the variation of surfaces by varying the coefficients of terms in several implicit equations, including in Equation 2. When implementing Equation 2 the authors fixed t = 0 and applied the relationship B = 1 − A to reduce the number of parameters to one. They then varied A from zero to one to study the changes to the surface topology. Applying the same relationship to our 3D design space, we once again considering the multiplicative property of equality to show that we can scale any values of A, B, and t by a common factor such that A and B satisfy the relationship B = 1 − A without changing the geometry of the lattice. As a result, the design space can be reduced to a 2D space with A (or B) and t.

As a result, the design space can be reduced to a 2D space with A (or B) and t. However, for reasons discussed in Section 3.1, we found that giving the optimizer independent control over A and B was preferable so we maintain three dimensions in the design space of our study.2.2 Definition and Evaluation of ObjectivesFor a given volume fraction, we seek to produce the stiffest version of the lattice possible. To this end, we consider two objectives. First, maximize the stiffness in the direction of compression (E11), which we redefine as: minimize the reciprocal of the stiffness in the direction of compression. Second, minimize the volume fraction (VF). These two objectives will produce a Pareto frontier of designs each of which, within the limits of the convergence criteria (discussed in Section 2.4), will be the stiffest for that volume fraction.

These two objectives will produce a Pareto frontier of designs each of which, within the limits of the convergence criteria (discussed in Section 2.4), will be the stiffest for that volume fraction. We apply two additional inequality constraints to the problem on the volume fraction objective, limiting the volume fraction to the range of [0.1, 0.9]. This lower limit avoids thin features that fail to mesh well, and this upper limit was imposed to prevent excessive evaluation time spent on designs that are not of particular interest to the authors. We now formulate the optimization problem as:min A,B,t s.t. Ω(1/E11, VF)0.1 ≤ VF ≤ 0.9To evaluate the first objective function, we use a linear finite element model using a voxel meshing approach based on Dong, Tang, and Zhao [21]. This is a computationally efficient ap- proach that can be used to generate a highly regular volume mesh directly from an implicit equation without first extracting a surface mesh from the field.

This is a computationally efficient ap- proach that can be used to generate a highly regular volume mesh directly from an implicit equation without first extracting a surface mesh from the field. The voxel mesh is used to generate a mesh of constant strain three dimensional eight node (C3D8) brick elements. The Young’s modulus (E) of each element is linearly weighted based on the number of nodes (n) of that element that are inside the lattice using Eq. 12.Eelement = Ematerialn 8For the optimization, we aim to simulate the far field response of the lattice, not that of only a single, isolated unit cell [5]. When performing physical testing of lattices, a minimum of 64 cells1961(11)(12)(in a 4x4x4 grid) is required to effectively eliminate edge effects [28]. However, in linear elastic simulation, we can take advantage of the periodic nature of the unit cell, by applying additional boundary conditions, to reduce the size of the model to a single unit cell of the lattice.

The IWP TPxS has a cubic crystalline structure (like the BCC), which is a subset of orthotropic materials that, in addition to containing three orthogonal planes of symmetry [21], are invariant to 90° rotations of the lattice. Thus, we would require only three simulations to compute the three values needed to generate the homogenized constitutive matrix. Because we are only interested in minimizing the compliance of the lattice in one direction, the constitutive matrix is not expressly required. We therefore further reduce the computational load by evaluating the response of the lattice in only one load case which yields the first column of the constitutive matrix C.

We therefore further reduce the computational load by evaluating the response of the lattice in only one load case which yields the first column of the constitutive matrix C. Because we do not calculate the full matrix, we cannot invert it to find the stiffness matrix, and we instead calculate an effective normalized stiffness in the direction of the compression load using Hooke’s law solved for the elastic modulus given in Equation 13 and use in the first objective in Equation 11:E11eff =σ11 dεC11where σ11 is the (1,1) element of the stress tensor, dε is the applied displacement, and C11 is the (1,1) element of the constitutive matrix. We calculate the reciprocal of E11eff to get our objective function, which we note is an analog for compliance as it was not generated using the full matrix.For the IWP lattices, the faces of the unit cell, shown in Figure 1b, are also planes of symmetry.

To account for the additional cells of the lattice that are not modeled, the nodes on the face of the cubic unit cell are restricted to move in the plane of the given face, to enforce symmetry, with the exception of nodes on the x = Px face which have a fixed, uniform displacement in the −x direction. We use an isotropic material model with Young’s modulus and Poisson’s ratio of 2 GPa and 0.394, respectively.When validating the results of the optimization and comparing against other lattice designs that share the same topology, the boundary conditions and performance metric are different from those described above.

The validation model setup is presented in Section 2.5 and both the optimization and validation model setups are summarized in Table 1.We note that because of the form of Equations 1 and 2, and the resulting rotational symmetry of the cubic crystal unit cell of the lattice, the response to a given loading in the y or z direction will always be the same as in the x direction. While this likely limits how much improvement in the x direction we can achieved, it results in a lattice that is more useful in application.

While this likely limits how much improvement in the x direction we can achieved, it results in a lattice that is more useful in application. Future work can explore the effects of allowing direction changes to the lattice geometry.To evaluate the second objective, we compute the value of Equation 2 over a 3D grid in x, y, z space with 200 points in each direction, The number of points inside the lattice (points where Equation 2 ≤ 0) is divided by the total number of points in the unit cell to calculate the volume fraction.1962(13)Table 1: Summary of boundary conditions applied to FEA models for optimization runs and validation runs that are introduced in Sections 2.2 and 2.5, respectively. The two models use different boundary conditions at the x = Px surface of the unit cell as well as different criteria to measure the performance of the model due to differences in software capability.

We use ∆ to signify the prescribed displacement in the stated direction.Boundary ConditionsBoundary Optimization RunsValidation Runs ∆x = 0 ∆y = 0 ∆z = 0 Fixed load in −x direction distributed across surface ∆y = 0 ∆z = 0 Model Performance Criteria∆x = 0 ∆y = 0 ∆z = 0x = 0 y = 0 z = 0∆x = −0.01Pxx = Px∆y = 0 ∆z = 0y = Py z = Pz1/E11Maximum nodal displacementPer RQ1.1, the designs on the Pareto frontier will be related to one another in the design space such that the coefficients A and B can be defined as functions of t, allowing us to produce an equation with a single input parameter, similar to other TPMS-based lattice designs; however, the voxel approach we use to evaluate the objective function results in discrete steps in the performance space, making the derivation of these relationships more challenging.2.3 Selection of Optimization AlgorithmIn this work, we use the Borg MOEA to solve our optimization problem defined in Eq. 11.

11. prior to applying Borg to this problem, two other optimization approaches were tested. First, we explored gradient based methods. These methods are compelling for this work because it was expected that they would result in a smooth Pareto frontier from which design space relationships could be derived. However, the voxel FEA approach used for expedient meshing and analysis does not pro- duce locally smooth gradients in the performance space. Because these methods rely on the slope of the gradient to converge, they could not be used. Next we explored purely genetic approaches, the results of these optimization runs indicated that these approaches were not appropriate, as some designs on the Pareto frontier failed to outperform the standard equation (which is within the design space). This lead to the exploration of MOEAs that could adapt to the problem, and the selection of Borg.

This lead to the exploration of MOEAs that could adapt to the problem, and the selection of Borg. Borg is a high performance adaptive algorithm that was developed to solve difficult multi and many objective optimization problems and was selected because its auto-adaptive search features resolved issues with tuning the other algorithms considered [29].2.4 ImplementationIn this section we briefly describe details pertaining to the implementation of Borg and the function it calls for evaluation of the objectives. When Borg makes a call, it first evaluates the feasibility of1963the design by computing the validity of each constraint and the distance from the constraint surface to the design, which is returned to Borg. If one or more constraints were violated, the analysis is ) bypassed, and the objective functions are returned to Borg as artificially high values , because Borg always expects real values to be returned.( i.e.

1020When Borg was called, we used a maximum number of iterations of 100,000 and ε values of 0.01 for both objectives. The epsilon values effect both the density of designs on the Pareto frontier as well as the stagnation/convergence criteria of the algorithm. Respectively these are because Borg utilizes an ε-box dominance archive and an ε-progress method. Our epsilon values were selected because they produced a Pareto frontier of the desired density of approximately 1 design per 1% change in volume fraction.Borg is available natively in both C and Python. The C implementation is more computationally efficient and offers wrappers to interface with other languages.

The C implementation is more computationally efficient and offers wrappers to interface with other languages. The C implementation of Borg was compiled into a MATLAB executable allowing it to be called within MATLAB, allowing Borg to interface with existing meshing and analysis codes written natively in MATLAB.The optimization was performed using MATLAB 2022a on a desktop workstation with an Intel Xeon W-2133 CPU and 128GB of RAM running Windows 10. Borg was run in series and the resources of the computer were not fully utilized, motivating parallel evaluation in future studies.2.5 Validation ModelsIn order to evaluate the performance of our optimized lattice in comparison to existing designs in the literature, we first developed matching models to represent each using nTopology’s internal analysis package.

nTopology is able to natively model, mesh, and analyze lattices based on implicit equations, struts, and high order Bézier curves, allowing for easier integration and analysis which are used to model the IWP TPxS, BCC, and optimized BCC from Lee, Zhang, and Gu [18] respec- tively. Like other FEA tools, CAD models can also be imported for evaluation of the optimized BCC lattices from Wang et al. [8] and Zhao et al. [19].The FEA was carried out using linear tetrahedral mesh elements that were generated by dis- cretizing the surface of a given lattice model with triangles before populating the volume of the model uniformly with tetrahedrons. The unit cells were modeled as 10mm cubes, and the tetrahe- drons were defined with an edge length between 0.1mm and 0.25mm depending on volume fraction and a fixed growth rate of 1.1 for all the models.

The size of the triangles in the surface mesh were matched to that of the tetrahedrons to produce consistent element sizes at the surfaces of the model.In addition to this change to the model setup, the validation models do not use the same perfor- mance metric. Instead, when validating the optimized structures against similar geometries from the literature, we use the maximum deflection of the model’s nodes to quantify stiffness instead of calculating E11 using Equation 13. Table 1 shows the model setups for both the optimization and validation cases. While the differences between the two models means the results cannot be compared against each other, they can be used to compare to other lattice topologies that are analyzed with the same boundary conditions.19640.10.20.30.40.50.60.70.80.9Volume Fraction020406080100120140Compliance AnalogStandard IWPOptimized IWPFigure 7: Borg output compared to standard IWP TPxS plotted in the performance space.

Showing that the optimized structures outperform the standard across the range of volume fractions considered.3 RESULTSWe present the Pareto frontier resulting from our Borg optimization in Figure 7. We also include the standard IWP TPxS in Figure 7, evaluated using the same model setup and solver to illustrate that the designs of the optimized version Pareto dominate those of the standard equation, answering RQ1. The optimization took 67 hours to converge within the limits given for the objectives, utilizing just over 85,000 evaluations, 72% of which fell within the feasible region.To quantify the improvement of the optimized lattice over the standard IWP TPxS across the Pareto frontier, we evaluate the standard IWP TPxS was at 200 points, equally spaced in t-space.

For each Pareto optimal design, the standard design from the dataset with the lowest volume fraction that was greater than that of the optimized volume fraction was used for comparison From this comparison the optimized design improves performance by between 2% and 18%. We tabulate these data in Appendix A for the 80 Pareto optimal IWP TPxS designs.3.1 Relationships in the Design SpaceWe now seek empirical relationships between A, B, and t for the Pareto optimal designs in the design space. In Figure 8, we plot A and B as functions of t and no relationship between A and t or B and t were found when considered independently. Both A and B vary across the range of t values with only a slight trend to the data. Initial curve fits that included all points shown in Figure 8 resulted in designs that performed worse than the standard IWP TPxS.

Initial curve fits that included all points shown in Figure 8 resulted in designs that performed worse than the standard IWP TPxS. 3D visualization19651-0.500.51 0.20.40.60.81A1-0.500.51t B=f(t) Pareto DesignsFigure 8: The A and B parameters of the designs on the Pareto frontier plotted as functions of the level set t. The designs in red are a subset of the Pareto frontier which are scaled and fit with a polynomial shown in Figure 10.of the design space in Figure 9 illustrated that a relationship between A, B, and t for some, but not all, of the points on the Pareto frontier was likely to exist. To derive a relationship, we again invoked the multiplicative property of equality and normalized the parameters of each design such that A = 2. After normalizing, in Equation 14 we plot the scaled B values against the scaled t values. In this scaled B–t space, the data follow a generally linear trend, the spread of which is significant.

After normalizing, in Equation 14 we plot the scaled B values against the scaled t values. In this scaled B–t space, the data follow a generally linear trend, the spread of which is significant. When fit, this results in under-performing designs. We note that in this space the designs exhibit a distinct upper bound that is of interest. To identify points on this edge, we apply the concept of Pareto optimally to the existing designs in this B–t space, with the objective of maximizing B and t. These points are marked in red in both Figures 8 and 10. We then fit the second-order polynomial presented in Equation 14, plotted in red in Figure 10.Bscaled = −0.007046t2scaled− 0.322tscaled + 1.972We now use the fixed value of A and the scaled B = f (t) relationship with Equation 2 to model lattice structures. We analyze the designs on this curve using the same model setup and solver in as the optimization, and in Figure 11 we plot them with the optimized and standard IWP TPxS designs.

We analyze the designs on this curve using the same model setup and solver in as the optimization, and in Figure 11 we plot them with the optimized and standard IWP TPxS designs. We find that the designs generated with the curve fitting are distinct from the Pareto optimal designs found using Borg. The curve fit designs Pareto dominate all of the standard designs, but do not Pareto dominate any of the Borg optimized designs, answering RQ1.1. We therefore provide two results from this work. First, 80 discrete designs that show improvement over the standard design with volume fractions ranging from 0.1 to 0.9 and second, a new single variable1966(14)Figure 9: Pareto optimal designs plotted in the design space. A trend that involves A, B, and t can be seen. However, a 3D surface fitting of would not yield single value functions for A and B for a given value of t.equation for modeling lattices which shows less improvement but is continuous.

However, a 3D surface fitting of would not yield single value functions for A and B for a given value of t.equation for modeling lattices which shows less improvement but is continuous. Both results have been packaged into custom nTopology blocks. These blocks make design and analysis of the lattices accessible and they can be downloaded from GitHub2 under a Creative Commons Attribution 4.0 International License.When applying this method to the Pareto frontier from the optimization run with two objectives we found that the points had no relationships that could be found and no significant differences be- tween the frontier presented in this work and the two objective Pareto frontier.

To avoid unnecessary confusion, we do not present the results of the two objective run.3.2 Comparison to Other LatticesIn addition to comparing the optimized and standard IWP TPxS using voxel based FEA in MAT- LAB that is shown in Figures 7 and 11, we compare the optimized designs against four additional lattices that share the same topology using nTopology’s FEA package. For this analysis, we use the Pareto optimal designs, not the sub-optimal fitting found in Section 3.1. We compare against the standard BCC lattice (shown in Figure 1a) as a baseline, and three optimized BCC lattices (shown2https://github.com/jwf23/Optimized_IWP_TPxS19675.5-5-4-3-2-10123t scaled Model fitting B=f(t)Figure 10: Down-selection of designs in the scaled B-t space for fitting B = f (t) using Pareto optimality and maximization of the scaled B and t values. Showing the polynomial fit to the Pareto front.in Figure 4) for a range of volume fractions.

Showing the polynomial fit to the Pareto front.in Figure 4) for a range of volume fractions. We discuss the definition of the FEA models used in Section 2.5 and summarized in Table 1.Visual comparisons between the different versions of the BCC and our optimized designs like those shown in Figures 12 and 13 highlight the difference in performance at low volume fractions, however at high volume fractions it visually unclear which version out performs and to what degree. To quantify the difference, we evaluate the standard BCC, Wang optimized BCC, and Zhao optimized BCC lattices at nine volume fractions matching volume fractions of the optimized IWP TPxS across the range of volume fractions. For the Lee optimized BCC, we compare to the nearest value from the optimized designs as both exist only at discrete volume fractions. These data as well as the percent difference between the compliance analog for each run and the optimized IWP TPxS are presented in Tables 2 and 3.

These data as well as the percent difference between the compliance analog for each run and the optimized IWP TPxS are presented in Tables 2 and 3. From the tabulated data, it is clear that the Pareto optimized IWP TPxS shows improvement over the other designs across the full range of volume fractions with one exception. At a volume fraction of 0.1089 it is outperformed by the optimized BCC with rounds between the struts from Wang et al. [8], addressing RQ2.The lattice in Zhao et al. [19] was optimized for a more limited range of volume fractions compared to this work. As a result, at larger volume fractions the performance is worse than the other designs, which can be seen in Table 2. At volume fractions of 0.4, 0.5, 0.6, 0.7, and 0.8, the Lee optimized BCC exhibits a change of topology caused by multiple minima along the Bézier curve that is used to define the struts [18]. This results in the struts intersecting away from the nodes of the lattice, ultimately changing the topology.

This results in the struts intersecting away from the nodes of the lattice, ultimately changing the topology. Because of this new topology, these lattices are not directly comparable to those in this study and have therefore been omitted. Table 3 contains1968Standard IWP0.10.20.30.40.50.60.70.80.9Volume FractionOptimized B=f(t)020406080100120140Compliance AnalogOptimized IWPFigure 11: Comparison between the B = f (t) designs, Pareto optimal, and standard IWP TPxS. The B- f(t) designs outperform the standard IWP TPxS.

The B- f(t) designs outperform the standard IWP TPxS. However, they fail to achieve the same performance as the optimized IWP TPxS.the three volume fractions: 0.1734, 0.2, and 0.3 for which the Lee optimized BCC exhibits the intended topology.3.3 Geometric Comparison to Standard IWP TPxSTo highlight spatial differences between the geometry of the standard and optimized IWP TPxS, we perform Boolean subtractions between the two at the same volume fractions (±1%) such that all material that is contained in one lattice but not the other can be identified. We show this comparison for five volume fractions in Figure 14. The color indicates which of the structures the different regions of “additional volume” are a part of. We see a trend in the optimized designs of material being drawn away from the “nodes” (at the corners and center of the cell) and moved toward the mid-span of the “beams”.

We see a trend in the optimized designs of material being drawn away from the “nodes” (at the corners and center of the cell) and moved toward the mid-span of the “beams”. This trend aligns intuitively with the initial FEA of the IWP TPxS, shown in Figure 1d, where the stress is minimal at the “nodes”. The one exception to the trend occurs at a volume fraction of 0.8925, and is shown in the last row of Figure 14. This is the largest volume fraction on the Pareto frontier and the only design where the relative weighting of the second term (coefficient B) of Equation 2 is lower than the standard in Equation 1.3.4 Homogenization and Anisotropy ComparisonAlthough the optimization of the IWP TPxS lattice we performed did not consider the response of the lattice in other directions or the anisotropy. Because in application the loading conditions will be more complex than uniaxial compression, understanding the directional response and anisotropy is useful.

Because in application the loading conditions will be more complex than uniaxial compression, understanding the directional response and anisotropy is useful. The IWP TPxS is highly anisotropic at low volume fractions [13]. To visualize and quantify1969Optimized IWP00.050.10.150.20.250.3Maximum DisplacementStandard BCC0.10.20.30.40.50.60.70.80.9Volume FractionFigure 12: Comparison of the Pareto optimal designs to the strut-based BCC lattice analyzed in nTopology. Showing that our optimized lattices outperforms the standard BCC across the range of volume fractions considered.the directional response and anisotropy, we can apply linear elastic homogenization based on the methods of Dong, Tang, and Zhao [21], and visualize the Young’s Modulus surface at volume fractions of 10.89%, 19.76%, and 29.81%, which we show in Figure 15.

From the elements of the homogenized constitutive matrices (Ci j), we compute the Zener ratio using Equation 15, which is a measure of the anisotropy of cubic crystals [30]:Z =2C44 C11 −C12Compared to the standard IWP TPxS at the same volume fraction, the optimization increases all non-zero elements of the constitutive matrices by the percentages shown in Table 4. From this, we find that the Young’s modulus increases in all directions, but the anisotropy also increases slightly in the optimized designs because the difference between the standard and optimized is not uniform for each element of the homogenized constitutive matrices.4 ConclusionsLattice structures in additive manufacturing offer a significant opportunity to reduce the weight of components without lose of structural integrity. The implementation of lattice structures can be further improved though the optimization of lattice structures, first at the unit cell level and then at the component level.

The implementation of lattice structures can be further improved though the optimization of lattice structures, first at the unit cell level and then at the component level. The workflow of the optimization of the IWP TPxS for a single load case has1970(15)Zhao Optimized BCC0.10.20.30.40.50.60.70.80.9Volume FractionWang Optimized BCCOptimized IWP00.050.10.150.20.250.30.35Maximum DisplacementFigure 13: Comparison of the Pareto optimal designs to the optimized BCC lattices generated by Wang et al. [8] and Zhao et al. [19] analyzed in nTopology. Showing that our optimized lattices outperform the optimized BCC designs with one exception at a volume fraction of 0.1089.been demonstrated and the resulting designs have been available as easy to use blocks importable to nTopology.In this work we have demonstrated an approach to the optimization of TPMS-based lattices for the IWP TPxS using the Borg MOEA to minimize compliance and volume fraction.

We share the results as blocks that can be imported into nTopology for lattice generation. In Section 1, we pose three research questions motivating the optimization of the IWP TPxS, and in Section 3, we present our results pertaining to these questions. We find that improvements over the standard IWP TPxS are possible across a wide range of volume fractions (0.1 to 0.9), addressing our first research question. The sub-question RQ1.1 relates to design space relationships between the Pareto optimal solutions. From the results of our optimization we derive a relationship between the coefficient B and the level set t after fixing A = 2. We demonstrated the outperformance of the standard equation; however, the performance of designs on this curve fall short compared to the discrete Pareto optimal designs from Borg. We developed FEA models based on the results of three optimizations of the BCC lattice, as well as the BCC as a baseline.

We developed FEA models based on the results of three optimizations of the BCC lattice, as well as the BCC as a baseline. The optimized IWP TPxS outperformed the four variations of the BCC with the exception of the design from Wang et al. [8] which out performed the optimized IWP TPxS at a single volume fraction of 0.1089. Therefore, we find that the optimized IWP TPxS outperforms the BCC designs for most volume fractions tested, but not all (answering RQ 2).Given the higher stiffness independent of direction, that is shown in Section 3.4, the Pareto optimal lattice replace the standard IWP TPxS in applications where higher stiffness for a given1971(a)(b)(c)(d)Figure 14: Differences between the optimized IWP TPxS (column a) and standard IWP TPxS (column d) at volume fractions of 0.1089, 0.2460, 0.4987, 0.7481, and 0.8925, from top to bottom. Column b shows the material that is part of the optimized lattice and not part of the standard, and column c shows the opposite.

Column b shows the material that is part of the optimized lattice and not part of the standard, and column c shows the opposite. From top to bottom, the percent differences by volume between the two lattices are: 14.6%, 2.2%, 8.2%, 4.8%, and 0.2%. Errors in rendering of thin sections are seen in columns b and c.weight is desired and in applications where smooth grading of the lattice is needed. The sub-optimal B = f (t) designs can be used in the same manner as the standard IWP TPxS.1972(a)(b)minmaxFigure 15: Young’s modulus surfaces for (a) the standard IWP TPxS and (b) the optimized IWP TPxS at volume fractions of 0.1089, 0.1976, and 0.2981 from left to right. The respective Zener ratios are: 10.4 and 11.3 (left), 4.84 and 5.15 (middle), and 2.85 and 2.89 at (right).

The respective Zener ratios are: 10.4 and 11.3 (left), 4.84 and 5.15 (middle), and 2.85 and 2.89 at (right). The range of the color bar is normalized for each volume fraction (column) to highlight the differences between the two designs.We propose four extensions of this work to expand the impact of our results. First, the blending of neighboring unit cells of different, discrete volume fractions should be explored, similar to work performed by Maskery et al. [5]. This technique would offer the ability to use multiple Pareto optimal volume fractions without sharp step changes in geometry at the interfaces. Second, our optimization can form the basis of a larger optimization scheme, where the unit cell is optimized for multiple cases and the results either balanced to produce a globally “optimal” unit cell or used in a component-scale optimization that utilizes optimal unit cells from different loading conditions depending on the local response of the component.

Third, application of this optimization method to other TPMS-based lattices including the TPnS and TPSf of the IWP as well as those based on other TPMS. Fourth, application of this optimization approach to equations for the IWP and other TPMS with additional coefficients that allow directional changes to the lattice.1973AcknowledgementsThis material is based on work funded by the Defense Threat Reduction Agency (DTRA) through the Naval Sea Systems Command (NAVSEA) under Contract No. N000024-18-F-6401, Delivery Order 21F8353. The opinions, findings, conclusions, and recommendations expressed in this material are those of the authors and do not necessarily reflect the views of NAVSEA. Nor do they reflect the views of DTRA or collaborators at NAVSEA.Table 2: Comparison of designs from the Pareto frontier to the standard BCC and optimized BCC versions from Wang et al. [8] and Zhao et al.

proving the Thermal Stability of Somos™ 6110 Photopolymer PatternsSuresh Jayanthi, Bronson Hokuf and John Lawton DuPont Somos™ Materials Group Two Penn's Way, Suite 401 New Castle, DE 19720 Tel:302-328-5435; Fax:302-328-5693Abstract:With the increased emphasis on functional testing and tooling applications using stereolithographic rapid prototyping (RP) techniques, the need for RP materials with improved thermal stability is becoming more pronounced. Glass transition temperature and Heat deflection temperature are two commonly used measures to determine the thermal stability of a polymeric material.This paper shows the effect of different postcuring techniques on the heat deflection temperature (HDT) of DuPont Somos™ 6110 photopolymer.

In addition, the benefits of using heat transfer mediums which provide neutral buoyancy and thus minimize thermal distortions are discussed.Introduction:The epoxy based stereolithographic photopolymers possess good dimensional and mechanical properties. The good dimensional properties are mainly attributed to the inherently low shrinkage during photocuring compared to the earlier acrylate based photopolymers. However, a major limiting factor of the first generation of epoxy photopolymers was their lower photo sensitivity. The recently introduced second generation epoxy resin DuPont - Somos™ 6110 provides significant improvements in photo sensitivity, allowing increased throughput from the stereolithography (SL) systems.

Somos™ 6110 is used in the Helium Cadmium laser systems like SLA-250™ from 3D Systems Inc., USA and STEREOS DESKTOPTM from EOS Inc., Germany.With the improvements in speed and physical properties possible from epoxy materials, the focus has now turned towards functional applications of the SL patterns. The SL patterns are now being widely used in a range of applications such as silicone rubber molding, injection molding as well as investment casting. With the ever increasing use in these tooling applications, the need for greater thermal stability of these patterns has become crucial.393This study focuses on methods aimed at improving the thermal stability of Somos™ 6110. The two main parameters that were looked into are the glass transition temperature (Tg) and the heat deflection temperature (HDT).

The two main parameters that were looked into are the glass transition temperature (Tg) and the heat deflection temperature (HDT). These two parameters are related to the internal changes in the material.Using the standard measuring techniques these two parameters were determined for the Somos™ 6110 epoxy photopolymer under various thermal postcure conditions. The relationships between these two parameters were studied with the objective of enhancing the HDT. Although no extensive quantitative data on the dimensional variations with thermal cure are available at this stage, a few preliminary results are presented.Thermal Stability:in polymeric materials. The first mechanism is a reversible process representing the softening of a material with increasing temperature and is used to characterize the thermal stability. The second mechanism is the irreversible decomposition of the material due to heat.

The second mechanism is the irreversible decomposition of the material due to heat. In this study we are primarily concerned about the first mechanism.Two differenttemperature-dependent mechanisms existThe thermal stability of polymeric materials can be expressed either by a time limit (at some specified temperature) up to which the temperature or material remains useful. A material is stable if no noticeable change occurs over any length of time or if any changes that occur are reversible. Typically, the point of thermal instability is observed by noting the onset of phase transitions that take place within the material. Two types of transitions are commonly observed on heating or cooling a plastic - a first order and a second order transition[l].A first order phase transition is observed when the material properties change abruptly. A second order phase transition is much more gradual and can usually be related to the amorphous segments of the polymer.

A second order phase transition is much more gradual and can usually be related to the amorphous segments of the polymer. The most prominent of the several second order phase transitions is the glass transition. It is an intrinsic material property characterizing a reversible softening behavior. According to ASTM El142, glass transition is defined as, "the reversible change in an amorphous material or in amorphous regions of partially crystalline material, from (or to) a viscous or rubber condition to (or from) a hard and relatively brittle condition". The glass transition temperature (Tg) is usually taken as the approximate midpoint of the temperature range over which the second order transition takes place.Besides the occurrence of transitions, certain properties of the plastic are continuously changing with temperature. The vibration of the chain segments increases with temperature and this motion weakens the cohesive forces between the material the molecules.

The vibration of the chain segments increases with temperature and this motion weakens the cohesive forces between the material the molecules. Decreased cohesion results (compromising its structural stability) and correspondingly affects the physical properties. On the basis of practical experience, criteria have been established whichin softening of394determine the limit of structural stability under load and heat. One such criterion is the HDT. A brief review of Tg as well as HDT will be provided below which will help us better understand the experimental results.Glass Transition Temperature:Glass transition is an important parameter in deciding the applications of many crystalline and non-crystalline polymers. It represents the temperature ranges at which the polymer matrix transitions from the glassy state to the flexible state. Below the Tg, however, the unreacted groups may not be able to crosslink due to lack of matrix mobility.

Below the Tg, however, the unreacted groups may not be able to crosslink due to lack of matrix mobility. The polymer chain segments cannot move around -- they are frozen in place and the behavior of the material changes from ductile to brittle. Above the glass transition temperature there is enough mobility within the object matrix to allow segmental chain motion, and the chains can move more or less freely around. As the temperature increases this freedom continues to increase until the matrix approaches the "liquid" state (if the polymer is not crosslinked). Even lightly crosslinked (or thermoset) polymers will have a glass transition that involves the motion of the linear chain segments between the crosslinks.

Even lightly crosslinked (or thermoset) polymers will have a glass transition that involves the motion of the linear chain segments between the crosslinks. If the crosslink density becomes high enough the glass phase transition will disappear.Some of the main molecular parameters influencing the Tg are [2]:Chain stiffness - Internal plasticization and - Intra molecular forcesThe mobility of the polymer chains is primarily affected by the barrier to their rotation around the backbone carbon-carbon bonds within the polymer matrix. This in turn is determined by the size of the substituent groups on the carbon atoms. The larger the substituent or the more hindered the rotation because of multiple anchoring (Van der Wall's forces between the side groups on the chain), the higher the Tg.Inclusion of low molecular weight materials (as side chains on the carbon carbon bonds) which generally have lower Tg, plasticizes the polymer.

Therefore, the higher the concentration of low molecular weight side chains the lower the Tg of the material. The internal plasticization can either be due to inter-chain or intra chain plasticization. Inter-chain plasticization is caused by small molecules that solvate and allow the polymer chains to move and become flexible. Water, for instance, can act as a plasticizer lowering the Tg. Intra-chain plasticization occurs when the primary chain of the polymer has long side chains (which act as plasticizers or chain softeners).Crosslinking has the opposite effect of a decrease in molecular weight. Increasing crosslink density increases Tg correspondingly. Most of the highly crosslinked thermoset polymers are amorphous.395Several methods exist to experimentally determine the Tg. One of the most commonly used methods is differential scanning calorimetry (DSC).

One of the most commonly used methods is differential scanning calorimetry (DSC). It is based on the power compensated null balance principle, in which the energy absorbed or liberated by the specimen is exactly compensated by adding or subtracting an equivalent energy to a heater located in the sample holder [3]. As the sample material passes through the glass transition upon heating, its (apparent) specific heat abruptly changes due to an increased motion in the material. In practice, specimens of thin flat disks of the material, usually less than 0.6 mm thick, are encapsulated inside an aluminum pan (0.08 mm thick) with a lid crimped in place to ensure uniform heat transfer. Platinum resistance heaters and thermometers are used in the DSC to carry out temperature and energy measurements. Another commonly used method for determining Tg is differential mechanical analysis (DMA). Values obtained using DMA are typically higher than those obtained using DSC [4].

Another commonly used method for determining Tg is differential mechanical analysis (DMA). Values obtained using DMA are typically higher than those obtained using DSC [4]. In this study DSC was used.Heat Deflection Temperature:As mentioned before, the Tg is an intrinsic material property characterizing the reversible softening behavior of the material. Many techniques have been developed to determine, in a well-defined way, the maximum usable temperature of a material due to softening. Heat deflection temperature (HDT) measurement is one such practical technique.The HDT is the temperature of the material at which a certain deformation is induced under flexural load of the test specimen. The test provides a measure of the thermo-mechanical stability of the material. A rectangular specimen 5" x 1/2" x 1/8" is placed on two supports 4" apart, immersed in a heat transfer medium (oil bath) with a load applied at its center to give a maximum fiber stress of 66 psi.

A rectangular specimen 5" x 1/2" x 1/8" is placed on two supports 4" apart, immersed in a heat transfer medium (oil bath) with a load applied at its center to give a maximum fiber stress of 66 psi. Higher loads were not used because there is no advantage to using higher loads when measuring deflection temperature of the present-day plastics with present-day instruments l . With the bar immersed and the load applied, the temperature is raised at a rate of 2 °C/min. The temperature at which a deflection of 0.010 inches occurs is noted as the HDT under flexural load of the material. The experimental testing reported in this study was done according to the ASTM D 648 standard.Experimental Discussion:The eventual physical properties are greatly influenced by the postcuring conditions used. The most common postcure techniques employed in curing green patterns are ultraviolet (UV) and thermal cure.

The most common postcure techniques employed in curing green patterns are ultraviolet (UV) and thermal cure. The UV cure is intended to polymerize any uncured liquid trapped within the solid by irradiating the patterns1ASTM D 648 - 82 (Reapproved 1988) - Standard Test Method for Deflection Temperature of Plastics Under Flexural Load. Annual Book of ASTM Standards 1993. Section 8.396with broad-band UV light. Thermal cure, however, induces additional polymerization by increasing the mobility of the polymer chains allowing additional chemical reactions to occur. In this study several thermal postcuring techniques to improve the thermal properties were investigated.regarding dimensional distortions become important. At elevated temperatures due to the softening behavior of the material, distortions are likely to occur in unsupported regions under the influence of gravity.

At elevated temperatures due to the softening behavior of the material, distortions are likely to occur in unsupported regions under the influence of gravity. One way of reducing these thermally induced gravity distortions is to heat treat the specimens in a dense heat transfer medium like silicone oil. The silicone oil also acts as a uniform heat transfer medium. The silicone oil used in this study is a commercially available DOW CORNING® 550 fluid which has a specific gravity of 1.1.With thermal postcuring techniques, concernsBesides the gravity induced distortions mentioned above, thermal treatment can induce distortions due to swelling. Though this behavior has not been extensively investigated, a few preliminary experiments have been performed and the results are discussed later.Another factor which can influence the dimensional stability as well as the glass transitions is the humidity.

The Somos™ 6110 is designed to produce good dimensional characteristics at higher humidities. Studies are underway to study the effects of humidity on the glass transition and will be reported at a later date.The experimental plan is essentially divided into three components. First, the Tg is determined using the DSC approach. Second, the HDT is determined using the ASTM D648 standard by thermally postcuring the specimens in air at various temperatures. Third, the HDT is determined by thermally postcuring the specimens in a silicone oil medium at various temperatures.For the Tg determination, sample specimens 0.2 mm thick and 25 mm cross section, were imaged in Somos™ 6110 using a SLA-250™. The imaged specimens were cleaned with tripropylene glycol monomethyl ether (TPM) and isopropyl alcohol to remove the residual uncured liquid polymer. Thermal postcuring was done in a programmable convection oven.

Thermal postcuring was done in a programmable convection oven. The green parts are considered to be postcured at 25°C (ambient temperature) while the UV parts are considered to be postcured at 40°C (the UV chamber temperature). For the HDT determination, the specimens were made with dimensions as specified in the ASTM D 648 standard. The green HDT specimens were cleaned as mentioned above. Four HDT specimens were used at each postcure condition. The green Tg and HDT specimens were later postcured under conditions described in Table 1.The Tg specimens were later measured for the glass transition using a DSC apparatus. The specimens were initially chilled below 0 °C before heating. A standard 10°C/min. rate of temperature increase was used for all measurements with a maximum temperature of 100°C. The results from the DSC measurements397are as shown in Figure 1.

A standard 10°C/min. rate of temperature increase was used for all measurements with a maximum temperature of 100°C. The results from the DSC measurements397are as shown in Figure 1. The Tg is measured as the inflection point in the glass transition phase curve on a temperature versus specific heat graph. The Tg of the material ranges from 39°C and 44 °C. This seems to indicate that no significant changes in the phase transitions are taking place as a result of thermal postcuring. The variation of Tg with thermal postcuring in silicone oil has not been investigated. However, no significant differences may be observed.Table 1: Postcuring Conditions for Tg Determination#1 2 3 4 5 6 7Postcure Temp. 25°C 40°C 60°C 8DoC 10DoC 12DoC 15DoCDescription Green patterns (ambient curing) UV cured patterns (UV chamber temp.) D.5hr ramp up; 1hr at 6DoC; D.5hr ramp down. D.5hr ramp up; 1hr at 8DoC; D.5hr ramp down. D.5hr ramp up; 1hr at 1DDoC; D.5hr ramp down.

The HDT appears to improve with the postcure operations. Table 2 describes the postcure conditions of the specimens immersed in a silicone oil medium whose density was selected to provide neutral buoyancy for the part and thus reduce distortions during postcure. Four specimens were used at each postcure condition. The heat deflection temperature of the specimens postcured in silicone oil are shown in Figure 3. As shown in Figures 2 and 3, the HDT appears to improve398with the thermal postcuring. The HDT of the specimens postcured in silicone oil appears to be higher than the air postcured specimens. This may be because the specimens postcured in oil had a longer cure cycle and therefore were held at an elevated temperature for a longer duration. Longer and more gradual rate of increase in temperature appears to improve the HDT.

25°C 40°C 60°C 80°C 100°C 120°C lSO°CDescription Green patterns (ambient curing) UV cured patterns (UV chamber temp.) 2hr ramp up; Ihr at 60°C; 2hr ramp down. 2hr ramp up; Ihr at 80°C; 2hr ramp down. 2hr ramp up; lhr at 100°C; 2hr ramp down. 2hr ramp up; lhr at 120°C; 2hr ramp down. 2hr ramp up; lhr at 150°C; 2hr ramp down.In the thermal Postcuring of the SL patterns dimensional stability concerns are critical. Efforts are underway to fully characterize the nature of dimensional distortions induced during thermal postcuring. A few preliminary studies have been done to determine the changes in the dimensions of the specimens. Our results show that both UV as well as thermal postcuring increases the dimensions3991 I I1 1: 11 _:1 1160by 0.1 % to 0.13%. Though the thermal postcuring in silicone oil does seem to improve the dimensional the specimens showed lesser gravity distortions.

Though the thermal postcuring in silicone oil does seem to improve the dimensional the specimens showed lesser gravity distortions. However, specimens which were conditioned in a desiccant chamber at 0% relative humidity for two days and thermally postcured showed a drop in the dimensional increase. Figure 4 illustrates the this phenomena. This result seems to indicate that thermal postcuring in the absence of moisture helps maintain the dimensional stability.increase,Figure 3: Influence of Postcure Conditions on HDT for Somos™ 611080I II II .-........I II I60U ° .5 40~ Q ==20... --M-----f--tM ::; _______ L I; L I '_-''-~I.

..... ~ 0 :: :: 0.10% <lJ0.15%l:I:ll:I:ljoooo(.!-<lJ~ 5eo: ..... -= Q<lJ ~ ;.. <lJ ~0.05%0.00%--J J60 tf UVDesiccant, 150°C (AIR)Postcure Condition400160Conclusions:The Tg does not vary significantly with thermal postcuring.The HOT, however, shows a increase of 15 to 25°C under the same postcure conditions.The Tg is lower than the HOT of the material.Thermal postcuring in silicone oil improves the HOT over postcuring in air perhaps due to extended time at temperature.Better dimensional stability can be achieved by conditioning the parts at 0% RH before thermal postcuring.References:[1] Gerhard F.L. Ehlers, "Thermal Stability of Plastic Materials", Engineering Design for Plastics, Reinhold Publishing Corporation, New York 1964.[2] James E. Mark, Adi Eisenberg, William W. Graesky, Leo Mandelkern, Jack L.

Mark, Adi Eisenberg, William W. Graesky, Leo Mandelkern, Jack L. Koenig, "The Glassy State: and the Glass Transition", Physical Properties of Polymers, American Chemical Society, Washington D.C, 1984.[3] Harvey E. Bair, "Glass Transition Measurements by DSC," Assignment of the Glass Transition, ASTM STP 1249, R.J. Seyler, Ed., American Society of Testing and Materials, Philadelphia, 1994, pp.50-74.[4] Chartoff, R. P., Weissman, P. T., and Sircar, A., "The Application of Dynamic Mechanical Methods to Tg Determination in Polymers: An Overview," Assignment of the Glass Transition, ASTM STP 1249, R.J. Seyler, Ed., American Society of Testing and Materials, Philadelphia, 1994, pp.88-107.Acknowledgements:This presentation is a result of the efforts of many people in the Somos™ Materials Group. We would like to acknowledge the valuable assistance rendered to us by Mimi Keating and Al Colombo in the evaluation of the test specimens. We would like to thank Dr. Daniel Mickish and Dr.

We would like to acknowledge the valuable assistance rendered to us by Mimi Keating and Al Colombo in the evaluation of the test specimens. We would like to thank Dr. Daniel Mickish and Dr. Glen Thommes for their continuous support and advice.Inquiries:Inquiries regarding this study may be directed to Suresh Jayanthi at the addresslisted or bye-mail tojayantss@esvax.dnet.dupont.com.401402.

proving volumetric accuracy of AM parts using adaptive slicing of Octree based structureSantosh Kumar Malyala*, A. Manmadhachary, Ravi Kumar.YDepartment of Mechanical Engineering, National Institute of Technology Warangal, Telangana - 506004, INDIAAbstractIn Additive Manufacturing (AM) processes, the layer-by-layer fabrication of complex geometries may lead to stair casing and thus error resulting in volumetric inaccuracies in the model. Using thinner slices reduces the staircase error and improves part accuracy but there is a tradeoff between number of layers and the build time for manufacturing part. This paper presents a octree based structure to improve the accuracy as well as reduces the build time. In the current work, firstly converting STL file into a modified boundary octree data structure (MBODS) and then calculating the non-uniform slice thicknesses (adaptive slicing) from the octree representation.

This slice thickness at any height is computed from the AM machine parameters and the smallest octree size at that available height. After the computation of the variable slice thicknesses has been completed, the part is virtually manufactured and the part errors are calculated. The virtually manufactured part and physical models are inspected to evaluate the volumetric errors. This algorithm uses an octree approach to improve the volumetric accuracy. And build time for the two different case studies are also doneKeywords: Additive Manufacturing, volumetric accuracy, octree structure, STL file, adaptive slicing.IntroductionThe Beauty of Additive Manufacturing is to build any complex geometry which is available in system . The Basic principle of AM is to build part layer by layer till part finishes [1]. The major advantage of this technology over conventional process to build complex shapes easily.

The Basic principle of AM is to build part layer by layer till part finishes [1]. The major advantage of this technology over conventional process to build complex shapes easily. If the design is available in CAD system then almost 100 times physical part can be produced though AM. In the recent years the demand for AM fabricated models are growing drastically. The major challenges with this technology is stair case effect occurs due to layered manufacturing principle, Which effects the surface quality and build time of the fabricating model [2]. The need to overcome this problems gave opportunity to develop various slicing algorithms. In traditional uniform slicing algorithms the layer thickness remains constant for single STL file. The major purpose of this STL file is optimum layer thickness along with maximum possible build quality and minimum build time, which is not easy to attain through uniform slicing.

The major purpose of this STL file is optimum layer thickness along with maximum possible build quality and minimum build time, which is not easy to attain through uniform slicing. But in adaptiveslicing there is opportunity to have variable layer thickness for the STL file, where layer thickness can be modified according to the geometry of the part. Adaptive slicing models comes out with better quality with less build time compared to uniform sliced part [3]. The layer thickness is one of the major parameter which effects on part quality. If the layer thickness is small then stair casing effect will be minimum and vice versa in case of curved and complex geometry. The layer thickness is also inversely proportional to the build time. Whereas in adaptive slicing the layer thickness is automatically defined on the basis of part geometry.

The layer thickness is also inversely proportional to the build time. Whereas in adaptive slicing the layer thickness is automatically defined on the basis of part geometry. This adaptive slicing gives a benefit of improved part quality and reduced built time in case of complex geometries also [4].The primary objective of this paper is to compute AM part with variable layer thickness to satisfy desired requirements. N. Siraskar has developed a particular algorithm which adapts layer thickness depending on the geometry of the part and calculates the desired thickness. The major problem with these existing algorithms is huge time consumption and complex in design [5]. In most of the adaptive slicing algorithms, the slice thickness is determined on the basis of various parameters like upper limit of permissible Ra value or cusp height.

In most of the adaptive slicing algorithms, the slice thickness is determined on the basis of various parameters like upper limit of permissible Ra value or cusp height. The maximum allowable cusp height was introduced by Dolenc and Makela in 1994 [6], which is accepted by various researchers for various applications and they were successful in that some of them proposed slicing algorithms in which CAD model initially divides in to horizontal layers with allowable cusp height. Those layers do not satisfy this requirement will further divided in to finer layers of uniform thickness [7-10].In the Local adaptive slicing also the part is divided into layers on the basis of allowable cusp height as discussed earlier [11, 12]. In this local adaptive slicing the initially it identifies the individual features are present in the part and then the slicing is done individually for the each feature.

In this local adaptive slicing the initially it identifies the individual features are present in the part and then the slicing is done individually for the each feature. In most of the cases different features with same height will be present and those different geometries need different slice thickness to attain desired tolerance. Edge node in octree has been introduced, in which edge node is represented by singe edge rather than removing as gray, black or white. This work also consists of representing composed surfaces which consists biquadratic representation by extensive octree structure with help of edge, face and vertex nodes [13, 15].MethodologyIn the current study a new approach is used in AM process to minimize the stare case effect without compromising in part quality and build time. Desired outcome is achieved with help of Modified Boundary Octree Data Structure (MBODS). The steps followed to achieve desired accurate STL with appropriate volume is as follows:Fig 1.

Desired outcome is achieved with help of Modified Boundary Octree Data Structure (MBODS). The steps followed to achieve desired accurate STL with appropriate volume is as follows:Fig 1. Flow chart of algorithmFragmentation of the initial STL File to required MBODS : Here in this octree representation the basic idea is to find the minimum and maximum values for the three axis. For example these calculated values are stored as Xmin and Xmax for the X axis and it continues same for the Y and Z axis respectively. Generally the STL file representation consists of only surface information and the name octree itself indicates a number eight. In the MBODS the part geometry is fragmented in to eight parts in all the faces of the geometry. These fragments are further classified in to the category of gray, white or black nodes.

In the MBODS the part geometry is fragmented in to eight parts in all the faces of the geometry. These fragments are further classified in to the category of gray, white or black nodes. The formulas used in this fragmentation are as follows:Calculate xmax, xmin, ymax, ymin, zmax, and zmin from the STL vertices (1)Calculate the scalar quantity, ls = max ((Xmax- Xmin), (Ymax- Ymin),(Zmax- Zmin)) (2)Calculate the cube depth, N = ceil(log2 (ls/Tmin), where Tmin is the minimum slice thickness (0.1 mm considered, it is minimum slice thickness of SLS) (3)Calculate the size of the root cube is L = Tmin × 2N (4)Initializing the gray, white or black values: The color of the node is represented on the basis of the specific property or identity of the node. The assignment of the white color represents that there was the presence of geometry and no overlapping is present.

The assignment of the white color represents that there was the presence of geometry and no overlapping is present. The gray color represents that the geometry is present in particular fragment and also there was overlapping of the geometry. The representation of black color means that there was no geometry present in this area and it is an empty or support area.Further finer division gray node: The gray color node indicates that the overlapping of geometry is present and which needs to be further fragmentation. This further division of fragment is required till the overlapping issue is resolved and there are no more gray nodes present or when node value attains to the level of layer thickness (0.1mm).

This further division of fragment is required till the overlapping issue is resolved and there are no more gray nodes present or when node value attains to the level of layer thickness (0.1mm). Which means that in the final geometry there are no presence of gray nodes or gray node with less than or equal to the size of layer thickness.Octree calculation for the gray nodes: The STL is the representation of the boundaries and here even though it is in octree representation still it is a STL representation, hence there will be no information about the volume. So some separate technique is required to calculate the volume of the gray node. Here Ray casting algorithm is used to calculate the volume of the required gray node.

So some separate technique is required to calculate the volume of the gray node. Here Ray casting algorithm is used to calculate the volume of the required gray node. For this ray casting to calculate the gray nodes and divided in to finer squares with number of N×N matrix.Finding the final volume of the geometry: The final volume geometry is calculated on the basis of the representation of the white nodes in the entire geometry. All the final black nodes are exempted from the calculation in finding the volume.

All the final black nodes are exempted from the calculation in finding the volume. All the rays represented with help of ray casting method are used to find out the rays that which are present inside or outside of the cube and the threshold value of the cubes are also taken in to consideration to perform whether further fragmentation is required or not and final volume of the geometry is calculated using the below formula:cubetotoutcubeVhhVO(4)Here Ʃhttot is the total sum of the height of all the rays and Hout is the sum of the lengths of all the ray segments which lie outside the part.Build time calculation: Typically build time is means the time taken by machine to build the geometry. this was calculated on the basis how much time spend by laser (since SLS process is considered in this paper) on each layer to build the geometry and time taken to spread the powder for each layer. Here the time taken to spread the powder for each layer almost remains constant.

Here the time taken to spread the powder for each layer almost remains constant. In case of uniform slicing the layer thickness remains constant but in case of adaptive slicing layer thickness varies, thus build time will be reduced. Here the build timer and numberof layers are inversely proportional to each other. the build time taken for the entire geometry is calculated by using the following formulas:The time to build one layer is then Ti= Tp + Ai/(Vlaser ×D)(5)And the time to build the entire prototype is: Tb =T1+T2+T3……Tn ; where n is the layer number The machine specifications considered in this work as follows: Layer thickness of Tmin= 0.1mm; Layer thickness of Tmax = 0.2 mm speed of the laser Vlaser = 5 m/s; laser beam diameter D= 0.030 mm(6)ResultsHere in this paper with help of the above procedure found out the volume for the two different case studies with irregular in geometry as shown in Fig 2.

In the below two cases the overlapping of geometry is present and the results obtained with help of above procedure are compared with the uniform slice thickness of 0.1mm and 0.2mm. The comparison consists of the number of layers and the volumetric error when compared to the original CAD geometry.Case 1:Initially the volume for the below shown fig is calculated using the uniform slicing method with a layer thickness of 0.1mm and 0.2 mm. the part is divided by uniform slicing and the volume of the each layer is calculated to find the complete volume of the part and same procedure is done for both the layer thicknesses. In the next step the volume of each node is calculated and added to get the final volume with help of octree method and results are tabulated as below.10 9 8 7 6 5 4 3 2 1 07.40161.01268.99041.12824.53872.0667123Fig 2. CAD Model of the object 1 Fig 3.

CAD Model of the object Fig 5. Volumetric error & Build time Vs slice thicknessTable 2. The comparison results of CAD Model of the object 2Slicing MethodNumber of slices Volume of CAD file Volume of VM part % Volumetric error Build TimeAdaptive Slice Thickness by MBODS 115 1020 1006.368 1.3364 10.396 minUniform Slice Thickness0.1 mm0.2 mm161 1020 1005.864 1.3858 14.533 min81 1020 1003.357 1.6316 7.309 minConclusionWith this approach the boundary data structure variable slice thickness is computed. Using this information part is manufactured virtually. This two test cases, in which boundary data structure of the part is calculated and these parts are then virtually manufactured using the computed slice thickness also inspected for volumetric error. Build time is also calculated to evaluate the effect of adaptive slicing on it. It is observed that with small amount of volumetric error can be achieved by short build time.

Build time is also calculated to evaluate the effect of adaptive slicing on it. It is observed that with small amount of volumetric error can be achieved by short build time. In future this algorithm can be developed in such a way that it can extract the part data directly from CAD model. In this paper volume occupancy is used as a criteria for termination of division of cube. But in future other criteria can be used for cube termination and checking the accuracy of the AM part. Also to calculate volume of the part inside the cube is done by using ray tracing method.

But in future other criteria can be used for cube termination and checking the accuracy of the AM part. Also to calculate volume of the part inside the cube is done by using ray tracing method. In future this can also be changed with other more efficient technique.References[1] T T Wohler, Additive Manufacturing and 3D Printing State of the Industry, Annual WorldwideProgress Report, (2013).[2] Byun H S and Lee K H, Determination of the optimal build direction for different rapid prototyping processes using multi-criterion decision making, Robotics and Computer-Integrated Manufacturing (2006) 22: 69–80.[3] Pandey P, Reddy N and Dhande S, Slicing procedures in layered manufacturing, a review,Rapid Prototyping Journal (2003) 9: 274–288.[4] Sreeram P, Dutta D, Determination of Optimal Orientation Based on Variable SlicingThickness in Layered Manufacturing.

and Makela I, Slicing procedure for layered manufacturing techniques, ComputerAided Design (1994), 1:4-12.[7] Yun J, Lee S, Jung Y, Development of a gap searching program for automotive body assemblies based on a decomposition model representation, Advances in Engineering Software (2015) 81:7–16.[8] Ghariblu H., and Rahmati S, “New Process and Machine for Layered Manufacturing of MetalParts,” ASME Journal of Manufacturing Science and Engineering (2014) 136:41004.[9] Paul R, and Anand S, Optimal Part Orientation in Rapid Manufacturing Process for AchievingGeometric Tolerances, Journal of Manufacturing Systems (2011) 30: 214–222.[10] Panhalkar N, Paul R., and Anand S, Adaptive Layering in Additive Manufacturing using a k-dTree Approach, Proceedings of the 41st NAMRC 2013, June 10–14, Madison.[11] Zhou M. Y, Xi J.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIN SITU MELT POOL MONITORING AND THE CORRELATION TO PART DENSITY OF INCONEL® 718 FOR QUALITY ASSURANCE IN SELECTIVE LASER MELTINGD. Alberts1, D. Schwarze1, G. Witt2SLM Solutions Group AG, Roggenhorster Straße 9c, 23556 Lübeck, Germany University of Duisburg-Essen, Institute for Product Engineering, Chair of Manufacturing Technology, Lotharstraße 1, 47057 Duisburg, GermanyAbstractAdditive Manufacturing looks back on a history of about two decades and today SLM® technology keeps moving as an integral element in industrial production environments. Sensitive markets such as energy, medical or aerospace have the highest quality standards for complex, safety-related and highly stressed components which are to be met at competitive costs for each build job and single part.

In this context process monitoring is necessary for documentation, qualification and at the same time it is expected to be able to detect process anomalies during the process. In addition to surface roughness, part density which mostly depends on volume energy, changing with laser power, scan velocity etc., is a distinctive quality feature of every component. This paper presents a method for a real time melt pool monitoring system based on photodiodes and the correlation between thermal emission and part density of Inconel® 718 with respect to volume energy deviation.Keywords: In situ; real time; melt pool monitoring; Selective Laser Melting; part density; Inconel® 718.IntroductionOver the last twenty-five years, Additive Manufacturing (AM) has established itself beside conventional manufacturing technologies in industrial production environments.

It is now possible to use AM for small-batch manufacturing or even quantity production with good productivity, repeatability and reliability.With respect to productivity, multi-laser or rather multi-optic SLM® technology is available with up to four simultaneously working laser sources on one part or multiple parts in one layer for a whole build job [WS14]. Moreover, laser power is being continuously increased to higher power, with up to 1000 W now available, and different kind of beam profiles available to increase build rates [BSH11, SLM17].One of the biggest advantages of AM due to geometrical freedom is the capability to create special and complex elements like lattice or honeycomb structures as well as overhangs or undercuts [Reh10]. Selected parts have the potential due to a redesign to implement lightweight structures as well as optimized fluid properties or stress reduction [RAH13, YHC16].

Selected parts have the potential due to a redesign to implement lightweight structures as well as optimized fluid properties or stress reduction [RAH13, YHC16]. On one hand, there is the capability to integrate a number of separate parts in a single unit and to eliminate further process iterations. On the other hand, there are other selected examples of reducing production steps by substituting conventional technologies with SLM® technology for example in existing repair applications [Sch14, Kro15].In addition to be able to produce nearly fully dense metal materials, mechanical properties can be comparable to conventional wrought or cast metals. However, multiple process parameters, machines and environmental conditions can make the final part quality.

However, multiple process parameters, machines and environmental conditions can make the final part quality. Significant time and cost is invested in non-destructive testing to analyses part quality and to ensure the quality requirements of the final component.it challengingto achieve repeatabilityinMotivationBesides the quality check of a final component after the job as part of post-processing it is required to reach further process understanding being able to follow the job results down to single part behavior. Process monitoring systems in the AM environment should have the capability to detect process anomalies, as well as provide a documented record of the entire process for traceability. The overall motivation is establishing a1481method to correlate signal or ratio response of process monitoring systems during the job and tracing them back to build part quality.

The overall motivation is establishing a1481method to correlate signal or ratio response of process monitoring systems during the job and tracing them back to build part quality. The goal is to differentiate afterwards (offline) or even during the build (online) between components, determining which reach their quality threshold and which not. An offline pass/fail job report would reduce time and cost for post-processing. Part-related real time monitoring and online detection of failed parts, which could be deleted from the current job, leads directly to a fabrication time reduction.

Part-related real time monitoring and online detection of failed parts, which could be deleted from the current job, leads directly to a fabrication time reduction. However, the procedure should end up with more stable part quality, reduction or prevention of quality variations and ensuring the reproducibility of ready-to-use AM parts [WPH14, Alb16, WSP14].State of the artIndustry and science are focusing on the development, application and usage of monitoring systems, which can divided due to the observation of laser or primary emissions as well as secondary emissions. Observation of secondary emissions means from the process emitted thermal emissions, imaging, ultrasound, low coherence interferometry (LCI) or acoustic procedures. Furthermore, we can distinguish between sensor layouts of off-axis or on-axis systems.One of the first approaches used acoustic measurement, which is based on measuring the surface acoustic wave velocity.

Smith et.al presented comparable results by electron microscopy and x-ray computed tomography relating to defect measurement like pores and roughness by changing the laser power between 70 W up to 200 W [SHP16]. There is actually no applied system directly to SLM technology but it look like a promising and visible possibility due to off-axis or on-axis observation in the future.Neef et al. integrated an on-axis sensor setup in a SLM® machine based on LCI. The system is able to measure topology part information during and between the recoating procedures in which anomalies based on the relation of powder and SLM® structures has been shown. [Nee14]Rieder et al. was using ultrasonic as an off-axis online as well as an offline inspection for test specimens out of Inconel® 718. Due to the online monitoring procedure and comparison of a CT-image the correlation of ultrasonic signals and the microstructure behavior with respect to laser power variations were shown [RSB16].

Due to the online monitoring procedure and comparison of a CT-image the correlation of ultrasonic signals and the microstructure behavior with respect to laser power variations were shown [RSB16]. Off-axis imaging is mostly and already commercial available to detect insufficient powder recoating [KAS15]. There can be an image documentation after powder application as well as after exposure. Based on an automatic analysis of the captured image an error detection is implemented with respect to manual adjustment of conditions and error detection limits. It supports immediately individual process errors due to powder supply and it controls post coating procedure. More in detail, Kleszczynski et al. presents a high-resolution CCD camera to evaluate solidified scan pattern after powder application layer by layer [KZS12].

More in detail, Kleszczynski et al. presents a high-resolution CCD camera to evaluate solidified scan pattern after powder application layer by layer [KZS12]. By means of the resulting layer topography, variations of volume energy are detectable and recoating errors were visualized based on overhang structures [KZS14].Another image-based off-axis process monitoring approach has been developed by Krauss using thermography. Krauss differs in his research between global inspection, layer-based inspection and local inspection (layer resolved observation). Via generating indicators, an adjusted analysis was used to create a performance measurement system, which enabled quality evaluation. His research focuses on system resolution, ability of detection and temperature determination. Different kinds of process phenomena were carried out which could correlate to build part quality. [Kra17]Other on-axis approaches used to extract thermal emissions via photodiodes. Pavlov et al.

Different kinds of process phenomena were carried out which could correlate to build part quality. [Kra17]Other on-axis approaches used to extract thermal emissions via photodiodes. Pavlov et al. use an f-theta lens as focusing unit and two photodiodes which are sensitive in the dedicate wavelengths of 1200 nm and 1400 nm. The relation between signal response and variations of layer thickness as well as hatch distance were shown with respect to single scan vectors. Afterwards, different scan strategies were used on test samples to analyze signal fluctuations and their correlation to melt pool instability. [PDS10]Berumen et al. describe an f-theta machine setup equipped with one photodiode and a spectral range of 780 nm up to 900 nm. Overhang structures, sharp contours as well as unplanned layer thickness variations were identified by means of process errors during the delivery of the build platform.

Overhang structures, sharp contours as well as unplanned layer thickness variations were identified by means of process errors during the delivery of the build platform. Due to heat accumulation, a correlation to part defects could be demonstrated for all three experiments with respect to photodiodes and their combination with a high-speed camera. [BBL10]1482Spears et al. were able to draw a fundamental relation between a resulting energy density via laser power variations as well as the correlation between the signal profile and the porosity. The machine setup as well as material and process parameters were not specified. [SG16]ObjectiveSurface roughness, hardness and part density are well-defined as distinctive global part quality features. Especially part density results of the sum with respect to part defects like pores, cracks, shrinkages etc. and leads to mechanical properties.

Especially part density results of the sum with respect to part defects like pores, cracks, shrinkages etc. and leads to mechanical properties. The Key Process Variables (KPVs) that affect the density of parts produced by the Selective Laser Melting process are laser power, scan velocity, hatch distance and layer thickness. These KPVs can be used to describe the “volume energy”, i.e. the total energy input into the component, in units of J/mm3. The target of this study is to show the correlation between thermal emission response, volume energy variation and final part density with a photodiode-based in situ melt pool monitoring system to enable the system to use as an online/offline quality assurance method for Selective Laser Melting.In this work, we investigate the photodiode response by changing the wall energy or volume energy for single as well as multiple scan vectors. Also, we assessed whether the part density correlates with the monitoring data.

Also, we assessed whether the part density correlates with the monitoring data. It is also an objective to analyze if the system is sensitive for process parameter change and if they are able to distinguish from each other.Experimental setupMachine and Optical SystemThe in situ process monitoring system to be developed requires the possibility of integration and usage in established manufacturing systems of SLM Solutions. In addition, the systems are designed to work regardless of machine sizes and machine types, e.g. multi-laser / multi-scanner systems and the kW power range. Figure 1 (a) shows an SLM280 system from SLM Solutions and the optic design of a modular-built manufacturing system as a part of the whole machine concept is shown schematically in Figure 1 (b).The fiber tip of a diode-pumped single-mode CW fiber laser is coupled into a collimation unit, which is adjusted for the particular machine system.

The collimated beam aligns axially to a 3D focusing unit, which allows a highly dynamic and high-precision positioning of laser focus along the optical axis. This unit is optimized for applications with high power densities in particular and replaces elaborate flat-field objectives. A scan system with digital control electronics and customized coating is used for high-precision positioning and deflection of the process beam to the particular process plane.The process radiation occurring from the exposure of several vector types and locally welded powder particles is a thermal emission signal, which diffusely emits from the melt pool. It is utilized as input value of the melt pool monitoring device. Beam splitter 1, shown in Figure 1 (b), reflects thermal emissions at the rear surface seen from direction of the fiber, which reflect back along the optical axis into the opposite direction of the beam path.

The upwards reflect wave lengths from primary optical axis is aligned vertically to the optical buildup of the melt pool monitoring device in which it will reflect on a further mirror through a focusing lens. Finally, beam splitter 2 splits the signals for photodiode 1 and photodiode 2.The front surface of beam splitter 1 transmits the laser wavelength and a proportion of the reflection servesas an input signal for photodiode 3 for a laser power monitoring device.1483Figure 1:(a) SLM 280; (b) Schematic assembly of optic design with melt pool monitoring and laser power monitoring.Melt Pool Monitoring and Laser Power Monitoring SystemThe melt pool monitoring system uses two photodiodes with different dedicated photosensitivity as a function of wavelengths – each is detecting near-infrared wavelengths, but the wavelength detected by ADC1 is different than the wavelength detected by ADC2.

Spectral response is permanently taken up by the individual photodiodes, forwarded to an ADC (analog digital converter) and provided in an FPGA (field programmable gate array). ADC1 is associated to photodiode 1 and ADC2 is associated to photodiode 2. Depending on machine and optical system setup as well as taking into account two different photodiodes the signal-to-noise ratio is at present up to 10.As described the reflected portion of the laser wavelength via beam splitter 1 is used as an input signal (measured laser power) of the present laser power (set laser power) by means of photodiode 3. The consumed electric current is converted into an electrical voltage and provided every 100 kHz via ADC1 and ADC2. The measured laser power is then transferred from the LPM device to the MPM device.

The consumed electric current is converted into an electrical voltage and provided every 100 kHz via ADC1 and ADC2. The measured laser power is then transferred from the LPM device to the MPM device. In addition to [ASW16] the two process monitoring systems are combined and both of them are directly linked to x/y-coordinates which is provided by an interface of the control device.Thermal emissions, the laser-on/off signal, the set laser power and the measured laser power is combined to the present x/y-coordinates (16-bit) as well as is provided from the FPGA to the control device every 10 µs. The FPGA controls the correct timing of the input signals, other potential operations and the subsequent measurement task of 10 µs. All data packages of a measurement task are stored every layer in a data file and can be traced back clearly due to time function and are able to read out. After completion of the exposure of a layer, a new file is created automatically.

After completion of the exposure of a layer, a new file is created automatically. During the process the output is represented in the form of a live 2D representation of the whole build platform, see figure 2 (a). The associated raw signals of a selected scan vector are shown in figure 2 (b).1484Figure 2:(a) 2D representation of thermal emissions (b) and complementary “Track Plot” of selected scan vectors.Experimental description, results and discussionThe following section describes three experiments, which build upon one another. Each experiment contains a description, procedure, results and discussion, which is evaluated independently from each other.

Each experiment contains a description, procedure, results and discussion, which is evaluated independently from each other. Based on the following parameter variations with respect to laser power (PL), layer thickness (DL), scan velocity (VS) and hatch distance (dT) the following energy calculations for the experiments can be made.𝐸𝑊 =𝑃𝐿 𝐷𝐿 × 𝑉𝑆[𝐽 𝑚𝑚2⁄](1)𝐸𝑉 =𝑃𝐿 𝐷𝐿 × 𝑉𝑆 × 𝑑𝑇[𝐽 𝑚𝑚3⁄]The energy of the single scan vector approach can be defined as the energy input of a wall, EW, respectively a contour or a first track of a hatch pattern shown in equation (1). The basis for calculating the energy of a volume shown in equation (2). EV is used to describe multiple scan vector within the single scan vector approach with respect to hatch distance as well as the volume of test specimens.An already developed, proved and validated material file of Inconel® 718 serves as a basis for the following experiments.

Standardized wall and volume energies are used in which energy “1” belongs to the developed parameter set and results in average part density of 99,8 % which based on 50 µm layer thickness. The possibility to build high quality parts under controlled conditions is a basis requirement before changing various influencing parameters and conditions and use them for further data approaches and correlation to various indicators.Basis intensity behavior of different process and measuring situationsThe sum of single measurement points with respect to the MPM device is relating to a single layer information shown in Figure 3 (a). Single test sample or build parts result on one hand to constant process parameters, like scan velocity etc., for specific scan vector types, for example support structures, contours or hatch areas. On the other hand, there are clearly defined but variable process parameters, like rotation scan strategies or scanning sequences.

On the other hand, there are clearly defined but variable process parameters, like rotation scan strategies or scanning sequences. These parameters rely in combination to machine environment and process conditions as well as positioning of part to different process and measurement situations for the MPM device. It is the aim to make different situations visible via comparing photodiode response when each photodiode is accurate adjusted to the current laser position.Although the process parameters for any given scan vector may be nominally identical, there may be a significant difference in the emission detected by the Melt Pool Monitor due to the prior conditions of the local area being scanned.

A distinction is made between “No Exposure” situation, means detection of powder material, Figure 3 (b), as well as the measuring situation of a first single scan vector without any other neighboring solidified scan vectors in the detection area, “Track 1” Figure 3 (c), which can represents a contour track. 1485(2)Furthermore, there is the situation of a second scan track surrounded by powder material due to the MPM detection area, “Track 2” Figure 3 (d), in which one single scan vector will be expose next to the first solidified scan vector on a certain distance. Next situation, see “Track n+1” Figure 3 (e), represents for example a hatch exposure. The MPM device is detecting the emitted thermal emissions of a single scan vector, which is positioned next to multiple and solidified single scan vectors in the first half of the detection area. Multiple scan vectors are exposed immediately after each other. The second half of the detection area is covered with powder.

Multiple scan vectors are exposed immediately after each other. The second half of the detection area is covered with powder. The last process situation, “Track 2n+1” Figure 3 (f), represents a single scan vector surrounded by solidified scan vectors on both sides of the current measurement point in the detection area.Figure 3:Basis intensity behavior of different process and measuring situations.For the current approach, single scan vectors were exposed on the first layer of a base plate with a standardized wall energy due to layer thickness, laser power and scan velocity variations. With respect to hatch distance variations a standardized volume energy is used. Single scan vectors are exposed in the opposite direction of the laminar gas flow in the build chamber. For statistic validation, each process and measuring situation is repeated five times, so the results illustrate the average value as well as standard deviation of thermal emissions.

For statistic validation, each process and measuring situation is repeated five times, so the results illustrate the average value as well as standard deviation of thermal emissions. The averaged value for every single scan vector is determined via automatically read out from the “Track Plot”, Figure 2 (b), of the MPM device, which has a measurement rate of 10 µs.Figure 4 shows a continuous decrease of photodiode response with respect to thermal emissions due to “Track 1”, “Track 2”, “Track n+1” and “Track 2n+1” independent of ADC1 and ADC2. The lower level of thermal emission response of about 5 % to 18 % in between each situation base on different conditions regarding to the current heat conduction situation. The surrounded powder of the “Track 1” approach acts as an isolator and leads the entered energy for heat dissipation limited to the base plate. During the “Track 2” approach, entered energy can additionally dissipate because of the neighboring scan vector.

During the “Track 2” approach, entered energy can additionally dissipate because of the neighboring scan vector. Even when there are already certain single scan vectors are placed next to each other, heat dissipation is apparently going on so rapidly that residual heat from neighboring scan vectors are not influencing the result during “Track n+1” as well as “Track 2n+1” approach. For every approach, thermal emission response for ADC2 is consistently 10 % to 14 % higher than ADC1. Standard deviations of ADC2 compare to ADC1 are in the first two approaches 21 % to 23 % higher than in the “Track n+1” or “Track 2n+1” approach. Within the two last approaches, standard deviations of ADC1 are 13 % to 18 % higher than ADC2.

Within the two last approaches, standard deviations of ADC1 are 13 % to 18 % higher than ADC2. One of the reasons for this fact could be a smaller number of spatters or ejected powder, which were recognized and could be detected from the MPM device in the experiment.148640003500] s t n u o C [ n o i s s i30002500ADC1m E2000ADC2la m r e h T15001000500Track 1Track 2Track n+1Track 2n+1Figure 4:Basis intensity behavior of different process and measuring situations.Single and multiple scan vectors wall and volume energiesIn the next experiment, it was used the single scan vector approach and made selective variations with respect to layer thickness, laser power, scan velocity and hatch distance. The purpose was to analyze the general relation between wall energy and volume energy variations and photodiode response.

The purpose was to analyze the general relation between wall energy and volume energy variations and photodiode response. The aim is to understand how different process parameters affect the MPM signals.As with the previous experiment, single scan vectors are exposed on the first layer of a base plate. For statistic validation, each process parameter is repeated five times, so the results illustrate the average value and standard deviation of thermal emissions. The averaged value for every single scan vector is determined via automatically read out from the “Track Plot” of the MPM device, which has a measurement rate of 10 µs.Energy variations with respect to layer thickness, laser power and scan velocity changes, lead to a standardized wall energy change as well as hatch distance changes to a standardized volume energy change. Beside the standard parameter “1”, there are six additional energy levels for each parameter, which change ±15 % up to maximum variations of ± 45 %.

Beside the standard parameter “1”, there are six additional energy levels for each parameter, which change ±15 % up to maximum variations of ± 45 %. Contrary to all other experiments, six application-oriented layer thicknesses are selected which correspond to a wall energy range of – 83 % to + 150 % compared to the standard parameter. Being able to produce all single scan vectors in one job, 10 mm long grooves relating to the selected layer thicknesses were milled into the base plate before.In Figure 5, photodiode response with respect to thermal emissions of ADC1 and ADC2 are shown over standardized wall energies due to layer thickness variations. The experiment simply increased or decreased the layer thickness from the standard value, and is presented in terms of the resulting wall energy, see equation 1. Increasing wall energy because of smaller layer thickness results in a decreasing photodiode response.

Increasing wall energy because of smaller layer thickness results in a decreasing photodiode response. With thinner layers (higher wall energy), heat is able to be conducted away quickly to the base plate, thus the thermal emission is reduced. The results can be divided into three sectors. Within a wall energy variation between “1” up to “2,50” no significant variation of thermal emissions can be noticed. From a lower limit of the layer thickness a balanced heat dissipation is given. Between “1” up to “0,33” there is a constant and almost linear increase of ADC1 and ADC2 which shows the range of measurement sensitivity. Within that range, wall energy changes correlates with photodiode response due to layer thickness variations. Sector three, between “0,33” up to “0,17” there is an almost constant photodiode response.

Sector three, between “0,33” up to “0,17” there is an almost constant photodiode response. From an upper limit of the layer thickness, the resulting powder quantity is insignificant, since only a small or no connection to the layer below or base plate exists.14874000] s t n u o C [ n o i s s i350030002500m E2000la m r e h T150010005000,170,330,501,001,672,50Standardized Wall EnergyFigure 5:Thermal emissions response over standardized wall energies due to different layer thicknesses.Figure 6 shows thermal emissions of ADC1, Figure 6 left, and ADC2, Figure 6 right, over wall energy variations by changing laser power and scan velocity as well as volume energy variations by changing hatch distance.

Average thermal emissions as well as standard deviation for each process parameter and their specific variation up to ±45 % shows same dimensions for ADC1 and ADC2 and can be used for evaluating the single and multiple scan vector approach at the same time.Laser power and scan velocity approach is linked to the “Track 1” experiment as well as hatch distance to the “Track n+1” experiment, shown in the prior approach “Basis intensity behavior of different process and measuring situations”.Thermal emission response of laser power and scan velocity variations correlates significantly with increasing wall energy. Despite identical wall and volume energy variations, there is a process-parameter- dependent photodiode response in which at lower wall and volume energy of 55 % with respect to the standard parameter “1” scan velocity and hatch distance is almost at the response level.

By energy increase there is a change of thermal emission response for laser power and scan velocity with an intersection at the standard parameter, which could allow a determination between process parameters due to thermal emission response.In this experimental procedure, hatch distance variations have actually no effect on thermal emission response, even not on a wide range of volume energy variation. One explanation could be within the single respectively multiple scan vector approach, that there is a direct heat conduction to the base plate. This procedure does not represent the fabrication process of a real part in which there is normally, depending on part dimension, not a comparable heat conduction as shown here.

However, correlation between hatch distance and the global quality feature of density is state of the art and the correlation to photodiode response should be checked on test specimens once again.1488ADC1ADC2A D C 1A D C 2Laser PowerScan VelocityHatch DistanceLaser PowerScan VelocityHatch Distance] s t n u o c [n o i s s i m El a m r e h T6500 6000 5500 5000 4500 4000 3500 3000 2500 2000 1500 1000] s t n u o c [n o i s s i m El a m r e h T6500 6000 5500 5000 4500 4000 3500 3000 2500 2000 1500 10000 , 5 5 0 , 7 0 0 , 8 5 1 , 0 0 1 , 1 5 1 , 3 0 1 , 4 5 Standardized Wall and Volume Energy0 , 5 5 0 , 7 0 0 , 8 5 1 , 0 0 1 , 1 5 1 , 3 0 1 , 4 5 Standardized Wall and Volume EnergyFigure 6:Thermal emissions response of single and multiple scan vectors over standardized wall and volume energies due to laser power, scan velocity and hatch distance variations.Correlation between photodiode response and part densityIn this third experiment, cylinders each with a height of 10 mm and 10 mm diameter, were produced in three different jobs as shown in Figure 7.

In the first two jobs the layer thickness was varied. In the third job, laser power, scan velocity and hatch distance were varied as shown at the bottom right in Figure 7.The first job was built with nine specimens for each layer thickness, starting with 40 µm. Afterwards, the next nine cylinders with a layer thickness of 30 µm was built on the same position and on top of the first existing specimens. The final nine cylinders had a layer thickness of 20 µm. The second job follows the same procedure as the first job with a layer thickness variation of 50 µm, 100 µm and 150 µm. Test specimens of the third job, based on 50 µm layer thickness, were built by changing volume energy in steps of 15 % each to a maximum of ±45 %.

Test specimens of the third job, based on 50 µm layer thickness, were built by changing volume energy in steps of 15 % each to a maximum of ±45 %. In fact, there are fifty-four specimens due to layer thickness variation and seven specimens each for laser power, scan velocity and hatch distance variations.For the current approach the MPM device had a measurement rate of 10 µs, which means a parameter depending quantity of measurement points between 5,6 million and 72,8 million per specimen.

The averaged value as well as standard deviation for every specimen is determined within a big data approach using computer language of Python.Density testing was performed for every specimen by doing metallographic studies via cutting, embedding, grinding, polishing and finally using standard software for density detection with light microscopy.1489Figure 7:Jobs and test specimens for process parameter variations.The results for the experiment, correlation between part density and photodiode response over volume energy with respect to variation of certain process parameter, can be seen in Figure 8 and Figure 9. The representation contains four process parameter variations with respect to layer thickness, laser power, scan velocity and hatch distance.Standardized volume energy “1” belongs to the standard parameter set as already mentioned and results in average part density of 99,8 % during the three jobs. Increases in volume energy above “1” also results in dense parts.

Increases in volume energy above “1” also results in dense parts. There are significant hot spots observed during manufacture of the specimens, which occur to higher surface roughness due to qualitative impression. This also leads to rising or decreasing thermal emission response for ADC1 and ADC2 between 10 % to 100%, depending on the process parameter.An increase in layer thickness decreases the current volume energy, and results in lower part density. Compare to all other process parameter variations in that paper due to a separated inspection for ADC1 and ADC2, a rising thermal emission results in significantly higher deviation of response due to ADC1 and ADC2 and resulting in denser parts. Results for layer thickness variations corresponds to contrary behavior and needs a different interpretation approach. For bigger layer thicknesses, powder is isolating most of the heat and results in significant hot spots as well as a lot of ejected powder and spatters.

For bigger layer thicknesses, powder is isolating most of the heat and results in significant hot spots as well as a lot of ejected powder and spatters. Smaller layer thicknesses result in a high level of heat conduction and lead to very shiny surfaces. That behavior can be seen in Figure 8 (left), which correlates inversely-proportional with part density. The inspection of ADC1 and ADC2 separated from each other contain among other melt pool size as well as melt pool temperature. The quotient evaluation of ADC1 and ADC2, see Figure 8 (right), eliminates the influence of the emitted area, the emissivity of the material, and contains the emissivity-independent temperature correlation. As seen, the emitted area has a significantly higher influence on the photodiode response and acts as an exponential influence compare to an almost linear characteristic of the melt pool temperature.

There it can be seen a clearly correlation between ADC1/ADC2 to part density and all results of analysis can be completely distinguish and classified from each other via a density deviation <0,1 %.1490Layer ThicknessLayer ThicknessADC1ADC2DensityADC1/ADC2Density] s t n u o C[n o i s s i m El a m r e h T5000 4500 4000 3500 3000 2500 2000 1500 1000 50010095908580]%[y t i s n e D2 C D A / 1 C D A0,70,610095908580750,5750,33 0,50 1,00 1,25 1,67 2,50 Standardized Volume Energy0,330,501,001,251,67Standardized Volume Energy2,50Figure 8:Correlation between photodiode response due to ADC1 and ADC2 (left) as well as ADC1/ADC2 (right) and volume energy variation via layer thickness change.Correlation between volume energy due to laser power as well as scan velocity variations and part density shows top left and right of Figure 9.

A dedicated inspection of ADC1 and ADC2, which is not shown here, indicates a close average value with respect to lower volume energy and rise up as the same time with higher volume energy. The increase of the independent photodiode response evaluation is almost linear up to a certain part density, which meets the reference volume energy level “1”. After that part density is higher than 99,7 %, the analysis of ADC1/ADC2 shows a clear correlation to part density as already shown in the layer thickness approach. Higher laser power correspond to a higher energy input per unit time and results in higher photodiode response. In this analysis, a resolution limit of the density can be specified with an accuracy of less than 0.1 %.At the bottom left in Figure 9 the correlation between volume energy variations depending on hatch distance and part density is shown. A smaller hatch distance, which means higher volume energy, leads to a rising part density of up to >99,9 %.

A smaller hatch distance, which means higher volume energy, leads to a rising part density of up to >99,9 %. A minimum part density of about 99,8 % is already reached by a smaller volume energy of -15 % due to the reference parameter. Contrary to previous investigation, hatch distance or volume energy variations, does not correlates with part density without any objection. The dedicated inspection of ADC1 and ADC2 as well as ADC1/ADC2 shows the linear characteristic up to a certain point. Although the photodiode response starts to decrease, part density is increasing up to the energy volume reference of >99,9 %. Following smaller hatch distances leads to a small but continuous decrease of part density of <0,1 %. The decrease of photodiode response at a volume energy offset of +45 % is in between the results of -30 % and -45 %, which belongs to a part density below 90 %.

The decrease of photodiode response at a volume energy offset of +45 % is in between the results of -30 % and -45 %, which belongs to a part density below 90 %. Bigger hatch distances belong to smaller volume energy and have no conductive connection to the vector next to the current one. Like in the layer thickness approach, just on a layer- based instead of a volume-based explanation, distance of two vector paths is so big, that the emitted area is expected to grow. On the other hand, melt pool temperature is obviously not immoderately rising. Bigger hatch distances indicate poorer heat dissipation and there is a well-defined limited area of emission because of the present solid material next to the current vector. Compared to the layer thickness approach, from the beginning of a certain layer thickness the emitted area is not limited due to the layer below. On the other hand, smaller hatch distances indicate better heat dissipation but a higher amount of residual heat.

The presented single and multiple scan vector approaches could show and indicate the sensitivities of the current monitoring system and enables the transfer of the method for real AM parts. Besides better process understanding these results will lead to further investigations of layer-based and local-based observations.A general photodiode response was presented by changing wall and volume energy with respect to single and multiple scan vectors via layer thickness, laser power, scan velocity and hatch distance variations. The correlation between parameter change and photodiode response for single scan vectors offers a first and fast basis of interpretation.

The correlation between parameter change and photodiode response for single scan vectors offers a first and fast basis of interpretation. It is possible to analyze if a certain process parameter or another influencing factor is able to detect as well as usable as a first investigation of the photodiode response resolution due to the parameter variation for the presented melt pool monitoring system.Finally, this study showed a correlation between part density and photodiode response due to layer thickness, laser power, scan velocity and hatch distance variations, which lead to a volume energy change. Independent and separate analysis of the two photodiodes offer further process understanding and lead to clear process relation behaviors in combination with the quotient of the two photodiodes.

Independent and separate analysis of the two photodiodes offer further process understanding and lead to clear process relation behaviors in combination with the quotient of the two photodiodes. The ratio response of1492photodiode 1 and photodiode 2 leads to a new quality assurance tool due to global quality diagnostic of density, which has been shown as an offline or post-processing approach. That method can be easily use as an online or in-line quality assurance tool. At that moment, due to photodiode response even a determination of the layer thickness is possible to distinguish to laser power, scan velocity and hatch distance. As widely known, independent from process parameter variation the volume energy has to have a certain critical minimum, which enables the fabrication of almost completely dense parts.

As widely known, independent from process parameter variation the volume energy has to have a certain critical minimum, which enables the fabrication of almost completely dense parts. In face of current challenges and objectives due to hatch distance variations it has been worked out, that a ratio response >0,66 corresponds to a minimum part density of 99,7 % that meets the demanding requirements of the industry.OutlookFuture works will concentrate on automatically, online “Global Quality Diagnostic” (GQD) due to any other global quality features like hardness or surface roughness. Within the online approach, there is the possibility to analyze global quality gradients as well as carrying out analyzes during the job. The GQD approach have to expand on various material systems. First results have to be statistically proven that there is the ability to define deviation limits and diagnostic resolutions as well as the documentation of a report.

 situ Printing – An Alternative Three Dimensional Laden Structure Fabrication MethodY. Liu1, W. Sun1, 2, 3, 41 Department of Mechanical Engineering, Drexel University, Philadelphia, PA, USA 2 Department of Mechanical Engineering, Biomanufacturing Center, Tsinghua University, Beijing, People’s Republic of China 3 Biomanufacturing and Rapid Forming Technology Key Laboratory of Beijing, Beijing, People’s Republic of China 4 Biomanufacturing Engineering Laboratory, Shenzhen Tsinghua Graduate School, Shenzhen, People’s Republic of ChinaEmail: weisun@drexel.edu and weisun@tsinghua.edu.cnAbstractRecapitulating a structure that mimics the anatomic geometries and intratissue cell distribution as in live organism is a major challenge of tissue engineering nowadays. Solid free- form fabrication (SFF) has been demonstrated as an efficient tool for this purpose.

Solid free- form fabrication (SFF) has been demonstrated as an efficient tool for this purpose. In this paper we presented a SFF based in situ printing method that is free of fabrication time frame and fabrication environment constrains. The fabrication parameters on strut formability, fabricated structural stability against gentle fluidic disturbance, and the integrity of the fabricated structure in cell culture environment were studied to assess the potential of the fabrication method on biomedical application. Based on the results, controlled strut formability can be achieved in an appropriate cross-linking deposition range. Alginate composition is the main parameter that dominates the stability and integrity of the fabricated structure.

Alginate composition is the main parameter that dominates the stability and integrity of the fabricated structure. A parameter set that can produce a stable scaffold with the ability to maintain its structure in cell culture environment for at least 15 days was optimized.IntroductionTissue engineering is an interdisciplinary realm that utilizes engineering and life science principles to advance the knowledge of tissue growth and behavior, which can be applied further toward the development of biological substitutes that restore, maintain, or improve tissue function. [1] Native tissue architectures are highly hierarchical in three-dimensional (3D) with heterogeneous cells and bioactive factors in precise disposition. This spacious allocation is found critical to tissue growth and function. Thus, it is a major challenge in tissue engineering nowadays to recapitulate the structures that mimic the anatomic geometries and intratissue cell distribution.

Thus, it is a major challenge in tissue engineering nowadays to recapitulate the structures that mimic the anatomic geometries and intratissue cell distribution. A famous pioneering work by Charles Vancanti and colleagues has demonstrated the therapeutic promise in which tissue-engineered cartilage was formed into human ear shape. [2] Similar cartilage structure was further extended to temporomandibular disc [3] and joint, [4] meniscus, [5] trachea, [6] intervertebral disc, [7] nasal tip, [8] and nasal septum, [9] as well as other bone structures as femoral shaft, [10] mandibular condyle [4] and distal phalanx. [11]While complex shapes were created, those technologies still lack the ability to incorporate cells and bioactive components to create a cellular spatial heterogeneities within the construct.

[11]While complex shapes were created, those technologies still lack the ability to incorporate cells and bioactive components to create a cellular spatial heterogeneities within the construct. To overcome this drawback, cell printing methods was introduced in tissue1635engineering to deliver an efficient method for fabricating 3D laden structure within which versatile distribution of heterogeneous cells can be achieved. To fabricate 3D laden structure incorporating bioactive components, various tools, such as inkjet [12, 13], laser-guided printing [14], two-photon polymerization and laser induced forward transfer [15], and 3D bioactive solid free-form fabrication [16-20], have been developed. Solid free-form fabrication (SFF) technology, often referred to as rapid prototyping, is analogous to 3D printing. In general, materials were fabricated and guided into different patterns for each layer, and subsequent layers were fused to build a complete free-form geometry.

In general, materials were fabricated and guided into different patterns for each layer, and subsequent layers were fused to build a complete free-form geometry. 3D bioactive SFF has been demonstrated high efficiency to build structure with complex geometries, as well to incorporate and maintain high viability of cells and activities of other bioactive component during the fabrication process.Various 3D bioactive SFF technologies have been explored to build 3D laden structures with cells encapsulated. Cohen et al developed a direct SFF technology where alginate was mixed with CaSO4 in advance as deposition material.[16] The pre cross-linked hydrogel was then mixed with cells and extruded for fabrication. Ahn et al modified this technology by introducing an aerosol cross-linking process.[17] In this method, alginate mixed with cells was pre cross-linked with CaCl2 and extruded for printing.

Ahn et al modified this technology by introducing an aerosol cross-linking process.[17] In this method, alginate mixed with cells was pre cross-linked with CaCl2 and extruded for printing. During the printing process, an additional CaCl2 solution was fumed with ultrasonic humidifier constantly flowing on the fabrication stage for further solidification. Khalil et al utilized an alternative method for cross-linking solidification.[18] They mixed the alginate solution with cells and extruded to reservoir where a CaCl2 solution with a depth same as the deposition layer thickness had been deposited for cross- linking. For each of the layer build up, CaCl2 solution was deposited in advance to elevate the solution level for another layer thickness. In addition to ionic cross-linking, temperature controlling system was also introduced to assist solidification.

In addition to ionic cross-linking, temperature controlling system was also introduced to assist solidification. Ahn et al developed a direct cell printing supplemented with low- temperature processing method.[19] In the dispensing process, alginate was mixed with cell-freezer solution and then mixed with cells as cell delivery material. The deposition material was extruded while a cooling system was circulating to cool the dispenser and working stage to solidify the extruded material. Fabricated scaffold was subsequently cross-linked with CaCl2. Zhao et al also utilized temperature controlled system for 3D cellular laden structure fabrication.[20] In their study, gelatin/alginate/fibrinogen hydrogel was used as cell delivery material and pre-mixed with cells. The deposition material was loaded to a temperature controlled deposition system where temperature was maintained as 25 ⁰C before extrusion and 10 ⁰C after extrusion.

The deposition material was loaded to a temperature controlled deposition system where temperature was maintained as 25 ⁰C before extrusion and 10 ⁰C after extrusion. This temperature difference dominates the solidifying process of the gelatin and thus delivers a porous scaffold structure with cells incorporated. Though the previous technologies have been demonstrated high cell viability and potential for cell proliferation, there are some limitations. The pre cross-linked alginate utilized a time dependent solidification process thus is constrained in a specific fabrication time frame. Reservoir crosslinking method is constrained by the topology of the reservoir and the scale size of the scaffold.

Reservoir crosslinking method is constrained by the topology of the reservoir and the scale size of the scaffold. Temperature control system requires an additional power supply and a specific working stage that limited the system’s portability.We presented an alternative fabrication method that is not constrained by a fabrication time frame, nor requires a specific printing stage or an additional power supply. A dual-nozzle system was utilized as deposition tool that mounted on the 3D motion system. In the dual-nozzle system, cell delivery material alginate and cross-linker material CaCl2 were loaded on each of the nozzle. The dual-nozzle extrusion tips were brought together for an in situ cross-linking process to fabricate 3D laden strut. Struts were designed and guided by the motion system to1636fabricate a complete 3D laden porous scaffold.

Struts were designed and guided by the motion system to1636fabricate a complete 3D laden porous scaffold. This paper reports on our preliminary study of in situ printing technology and the parametric characterization of the system on developing 3D porous laden structure. Deposition material was extruded and in situ cross-linked with variant fabrication parameters into struts which were guided to multiple layers stacked together as scaffold. The strut formability, fabricated structural stability against gentle fluidic disturbance, and the integrity of the fabricated structure in cell culture environment were studied to assess the potential of the fabrication method on biomedical application.Materials and MethodsCell-dispensing systemA proprietary 3D bio-printer was developed in this study to fabricate 3D laden structure. The 3D bio-printer includes a CAD/CAM integrated 3D motion system and a multi-nozzle biopolymer deposition system.

The 3D bio-printer includes a CAD/CAM integrated 3D motion system and a multi-nozzle biopolymer deposition system. The motion system was composed by three orthogonal robotic arms that move the loaded deposition nozzles in a 3D pre-defined pathway. The path can be programmed using an in- house software that is designed for CAD models. A CAD scaffold model that imported was first converted into STL format, and then sliced into multiple layers for toolpath generation. The path file was then imported to the motion system for freeform fabrication with variant motion parameters controllable. The motion resolution is 10 µm in the three orthogonal directions and has minimum velocity of 100 µm/ s.The multi-nozzle deposition system is mounted on the 3D motion system. Each nozzle deposition system consists of an air pressure supply, a material delivery tube, a dispensing nozzle system with pneumatic control, and a replaceable needle tip.

Each nozzle deposition system consists of an air pressure supply, a material delivery tube, a dispensing nozzle system with pneumatic control, and a replaceable needle tip. Deposition solutions were loaded in material delivery tube. The solution in material delivery tube is driven by a controllable air supply for flow rate control and continues with a dispensing nozzle system that is digitally controllable for turning on and off by a pneumatic control supply. The digital switch is integrated to PC and is coordinated to the motion system while 3D fabricating. The deposition solution goes through the dispensing nozzle system is finally extruded out through the interchangeable needle tip. Breathing air (Airgas) was used for air supply with high precision air pressure regulator for flow rate tuning. The in situ printing technology requires a dual nozzle system with the dispensing tip positioned together for immediate cross-linking.

The in situ printing technology requires a dual nozzle system with the dispensing tip positioned together for immediate cross-linking. The dual nozzle system was loaded with deposition material (alginate) and cross-link material (calcium chloride) respectively. During the fabrication, appropriate amount of solutions are extruded and quick cross-linking of the deposition material occurred to form the gel strut.1637Figure 1. 3D bio-printer system and multi-nozzle deposition systemChemical formulationAlginic acid sodium salt from brown algae with low viscosity and low glucuronic acid content (A0682, Sigma), and the product with medium viscosity and high glucuronic acid content (71238, Sigma) were selected as deposition material for this study. Those alginate products and their mixtures were dissolved in deionized (DI) water at various concentrations and left magnetic stirred overnight to ensure homogeneity and consistent viscosity.

Those alginate products and their mixtures were dissolved in deionized (DI) water at various concentrations and left magnetic stirred overnight to ensure homogeneity and consistent viscosity. An ionic cross- linking solution was prepared by dissolving calcium chloride (ACS grade anhydrous, amresco) in DI water at various concentrations. DPBS, 1X (Dulbecco’s Phosphate Buffered Saline) was used for scaffolds rinse.Strut formabilityThe strut formability of in situ printing was examined by measuring the extruded solidified strut size and comparing it with expected strut size. An analytic model was developed to determine the expected strut size.

An analytic model was developed to determine the expected strut size. The analytic model was developed based on the assumption that the alginate is cross-linked within a certain time frame when being extruded out thus deliver a cylindrical gel strut and limited swollen occurred, the diameter of the strut (D) can be expressed as a function of the fluid rate (Q), and the speed of the nozzle movement (v): D = ƒ(𝑄, 𝑣) = √4𝑄 𝜋𝑣. The fluid rate further is a function of the process parameters including theoperating pneumatic pressure, viscosity of the fluid, and the nozzle diameter.One can imagine that the strut formation includes two processes: deposition material extrusion, and extruded deposition material cross-linking. The deposition material extrusion process is described by the analytic model, thus the involved parameters are examined: deposition flow rate; nozzle size and nozzle travel speed.

The deposition material extrusion process is described by the analytic model, thus the involved parameters are examined: deposition flow rate; nozzle size and nozzle travel speed. The cross-linking process parameters are: alginate viscosity which regulates the ability of extruded material to hold the cylindrical shape during the cross-linking process; cross-linker concentration and cross-linker to deposition flow rate ratio those regulate the amount of cross-linker ions (calcium ions) available to alginate. During the examination of each parameter, the other parameters were set fixed while variant values of the examined parameter were tested to unveil the mere effect of the parameter on strut1638formability. Since the cross-linker concentration and cross-linker to deposition flow rate ratio are dependent parameters, they were examined as pair combination. The examined values of the parameters were listed in Table 1.

The examined values of the parameters were listed in Table 1. During the examination of the strut formability, the deposition material was selected as a mixture of the alginate product (1:3 as 71238: A0682). The strut size printed by each set of parameters was measured by optic microscope. Five samples for each set of parameter were examined. Mean strut size and standard deviation was calculated.Default valueExamined valuesDeposition parameters:Deposition flow rateNozzle size Nozzle travel speed Cross-linking parameters: Alginate concentration (viscosity) Calcium concentration Cross-linker to deposition flow rate ratio23.56 𝑚𝑚3/ 𝑚𝑖𝑛 250 𝑢𝑚 8 mm/s4.5% 5%3:122.00, 23.56 and 26.00 𝑚𝑚3/𝑚𝑖𝑛 200, 250 and 330 𝑢𝑚 4, 8 and 12 mm/s3%, 4.5% and 6% 1%, 3% and 5%1:1; 2:1; 3:1; 4:1; and 5:1Table 1. Values examined in the strut formability study. Default values were the values used when another parameter was examined.

Values examined in the strut formability study. Default values were the values used when another parameter was examined. Examined values were the set of values tested for the examining parameter.Structural stabilityBased on the result of the strut formability, the parametric values that succeeded to form the strut according to the analytic model were carried out for scaffold fabrication and stability examination. The stability study was designed to examine the capability of the scaffold to hold the structure against gentle fluidic disturbance.Besides the deposition parameters and cross-linking parameters, the material composition was also examined. Two types of alginate products and the mixture of those with variant ratio were studied. The alginates were medium viscosity with high glucuronic acid content (71238) versus low viscosity with low glucuronic acid content (A0682).

The alginates were medium viscosity with high glucuronic acid content (71238) versus low viscosity with low glucuronic acid content (A0682). It was found that fabricated hydrogel scaffold structure with low viscosity and glucuronic acid alginate are more stable and able to sustain the structure against normal fluidic disturbance. However, when saturated within cell culture medium, the hydrogel structures would be degraded quickly. Conversely, hydrogel scaffold structure with medium viscosity and high glucuronic acid content are more resistant to degradation when saturated in cell culture medium, but the structure are less stable and very easy to fall apart under mild fluidic disturbance. Thus, the examination of the mixture with both products was proposed to determine the optimal mixture ratio for biomedical application. The examined values of parameters were listed in Table 2.During the scaffold study, 3D hydrogel scaffolds were fabricated.

The examined values of parameters were listed in Table 2.During the scaffold study, 3D hydrogel scaffolds were fabricated. Parallel strut pattern of 15 mm × 15 mm with 750 um between struts was designed and 10 perpendicular laid out layers were stacked to build the 3D scaffold. The elevation of each stack layer was set to be 0.2 mm. The scaffolds were then moved to 35 mm petri dish with DPBS immersed. A gentle fluidic1639disturbance was introduced subsequently by orbital shaker of 100 rpm for 5 minutes.

The scaffolds were then moved to 35 mm petri dish with DPBS immersed. A gentle fluidic1639disturbance was introduced subsequently by orbital shaker of 100 rpm for 5 minutes. The structure before and after disturbance were compared for the examination of stability.Default valueExamined valuesDeposition parameters:Deposition flow rateNozzle size Nozzle travel speed Cross-linking parameters: Alginate concentration (viscosity) Calcium concentration Cross-linker to deposition flow rate ratio Material composition:23.56 𝑚𝑚3/ 𝑚𝑖𝑛 250 𝑢𝑚 8 mm/s4.5% 5%3:14.5% and 6% 3% and 5%2:1; 3:1; 4:1; and 5:1A0682: 712383:1Pure A0682; 3:1; 2:1; 1:1 and pure 71238Table 2. Values examined in the scaffold structurability and stability study. Default values were the values used when another parameter was examined.

Values examined in the scaffold structurability and stability study. Default values were the values used when another parameter was examined. Examined values were the set of values tested for the examining parameter.Structural integrityThe scaffolds that can sustain the structure from stability study were further examined for structural integrity study. The integrity study was intended to examine the capability of fabricated scaffold hold its structure in cell culture conditions. The scaffolds were immersed within cell culture medium (89% DMEM, 10% FBS, and 1% penicillin streptomycin) and incubated in 37 °C with 5% CO2. Values examined for this study was listed in Table 3.

The scaffolds were immersed within cell culture medium (89% DMEM, 10% FBS, and 1% penicillin streptomycin) and incubated in 37 °C with 5% CO2. Values examined for this study was listed in Table 3. The structure integrity was examined by checking the 3D structure at day 0, 1, 3, 5, 7, 9, 11, 13, 15.Default valueExamined valuesDeposition parameters:Deposition flow rateNozzle size Nozzle travel speed Cross-linking parameters: Alginate concentration (viscosity) Calcium concentration Cross-linker to deposition flow rate ratio Material composition: A0682: 7123823.56 𝑚𝑚3/ 𝑚𝑖𝑛 250 𝑢𝑚 8 mm/s4.5% 5%3:14.5% and 6% 3% and 5%2:1; 3:1; 4:1; and 5:13:1Pure A0682; 3:1Table 3. Values examined in the structural integrity study. Default values were the values used when another parameter was examined.

Values examined in the structural integrity study. Default values were the values used when another parameter was examined. Examined values were the set of values tested for the examining parameter.1640ResultStrut formabilityA set of parametric experimental studies were conducted to assess the effect of deposition parameters and cross-linking parameters on the strut formability. Fabricated struts were measured under optical microscope. Each data point was an average of five strut size fabricated with same set of parameters. The deposition parameters results were showed in Figure 2- 4, and the cross-linking parameters were showed in Figure 5, 6 and Table 4. Each of the experiment has one parameter examined to explore the effect of that parameter on strut formability and default values were used for the rest of the parameters.

Each of the experiment has one parameter examined to explore the effect of that parameter on strut formability and default values were used for the rest of the parameters. Since the cross-linker concentration and cross- linker to alginate flow rate ratio were dependent for cross-linker ions availability description, they were examined as pair combination.From the results of strut formability study on deposition parameters, the developed analytic model can be verified. In the condition when sufficient cross-linking can be achieved, the strut size is predictable by the analytic model. Variation of the deposition parameter value would result in strut size change. The precision of the strut size can be achieved within 10 um using the proprietary in situ fabrication system.The examination of cross-linking parameters showed that the strut formability is related to the cross-linking conditions.

When the formability was examined as a function of alginate concentration, it can be found that expected strut size can be formed when the alginate concentration is higher than 4.5%. In the examination of strut formability as a function of cross- linker availability (CaCl2 concentration and CaCl2 to alginate flow rate ratio), the results showed that 1% calcium chloride cannot form expected strut size regardless the flow rate ratio utilized. The expected strut size can be achieved when the cross-linker to alginate flow rate ratio is higher than 3 when using 3% calcium chloride solution, and 2 when using 5% calcium chloride solution.

The expected strut size can be achieved when the cross-linker to alginate flow rate ratio is higher than 3 when using 3% calcium chloride solution, and 2 when using 5% calcium chloride solution. As a summary of the cross-linker availability, strut formability of 4.5% concentration can be achieved within an appropriate cross-linking parameters range showed in figure 6D.Strut Formability in regarding to Deposition Flow Rate300.00m u n275.00ie z i St u r t S250.00225.00200.00212226 24 Deposition Flow Rate in mm3/ min2325271641Flow rate: 23.56 𝑚𝑚3/𝑚𝑖𝑛 Strut Size: 253 𝑢𝑚 Figure 2. Strut formability in regarding to deposition flow rate: the flow rate examined were 22.00 𝑚𝑚3/𝑚𝑖𝑛, 23.56 𝑚𝑚3/𝑚𝑖𝑛 and 26.00 𝑚𝑚3/𝑚𝑖𝑛.

Strut formability in regarding to deposition flow rate: the flow rate examined were 22.00 𝑚𝑚3/𝑚𝑖𝑛, 23.56 𝑚𝑚3/𝑚𝑖𝑛 and 26.00 𝑚𝑚3/𝑚𝑖𝑛 . The struts size were 219.40 𝑢𝑚 , 250.20 𝑢𝑚 and 280.40 𝑢𝑚 with standard error of 3.21 𝑢𝑚, 2.59 𝑢𝑚 and 4.10 𝑢𝑚 respectively.Flow rate: 22.00 𝑚𝑚3/𝑚𝑖𝑛 Strut Size: 216 𝑢𝑚Flow rate: 26.00 𝑚𝑚3/𝑚𝑖𝑛 Strut Size: 280 𝑢𝑚Strut Formability in regarding to Nozzle Size325.00m u n300.00ie z i S275.00t u r t S250.00225.00200.00150200250 Nozzle size in um300350Nozzle Size: 250 𝑢𝑚 Strut Size: 253 𝑢𝑚 Figure 3. Strut formability in regarding to nozzle size: the nozzle size examined were 200 𝑢𝑚, 250 𝑢𝑚 and 330 𝑢𝑚.

Strut formability in regarding to nozzle size: the nozzle size examined were 200 𝑢𝑚, 250 𝑢𝑚 and 330 𝑢𝑚. The strut size were 270.80 𝑢𝑚, 250.20 𝑢𝑚 and 300.00 𝑢𝑚 with standard error of 2.49 𝑢𝑚, 2.59 𝑢𝑚 and 2.55 𝑢𝑚 respectively.Nozzle Size: 200 𝑢𝑚 Strut Size: 269 𝑢𝑚Nozzle Size: 330 𝑢𝑚 Strut Size: 301 𝑢𝑚1642Strut Formability in regarding to Nozzle Travel Speed400.00m u nie z i St u r t S350.00300.00250.00200.00150.00100.000481216Nozzle Travel Speed in mm/sNozzle Travel Speed: 8 𝑚𝑚/𝑠 Strut Size: 253 𝑢𝑚 Figure 4. Strut formability in regarding to nozzle travel speed: the nozzle travel speed examined were 4 𝑚𝑚/s, 8 𝑚𝑚/𝑠 and 12 𝑚𝑚/𝑠.

Strut formability in regarding to nozzle travel speed: the nozzle travel speed examined were 4 𝑚𝑚/s, 8 𝑚𝑚/𝑠 and 12 𝑚𝑚/𝑠. The strut size were 372.80 𝑢𝑚, 250.20 𝑢𝑚 and 203.20 𝑢𝑚 with standard error of 7.05 𝑢𝑚, 2.59 𝑢𝑚 and 3.56 𝑢𝑚 respectively.Nozzle Travel Speed: 4 𝑚𝑚/𝑠 Strut Size: 371 𝑢𝑚Nozzle Travel Speed: 12 𝑚𝑚/𝑠 Strut Size: 202 𝑢𝑚Strut Formability in regarding to Alginate Concentration (Viscosity)300.0m u n275.0ie z i S250.0t u r t S225.0200.01.5%3.0%4.5%6.0%7.5%Alginate Concentration (w/w)1643Alginate Concentration: 4.5% Strut Size: 253 𝑢𝑚 Figure 5. Strut formability in regarding to alginate concentration (viscosity): the alginate concentration examined were 3.0%, 4.5% and 6.0% (w/w).

Strut size characterization as a function of CaCl2 to alginate flow rate ratio in different CaCl2 concentration (1%, 3% and 5% for A, B and C respectively). It can be summarized that the strut formability can be achieved within an appropriate cross-linking parameter range (shadow area in D).Cross-linker to alginate flow rate ratioCross-linker concentration1:1Mean SD Mean2:13:14:15:1SD Mean SD Mean SD Mean SD164461%391.8 5.22 327.8 10.47 289.4 4.16 280.0 9.41 285.8 6.063%5%294.6 5.50 277.4 5.32 253.0 4.06 251.8 2.95 252.8 4.02 283.4 4.04 247.6 5.50 248.6 3.58 250.2 4.97 248.2 3.96Table 4. Strut formability in regarding to cross-linker availability (CaCl2 concentration and CaCl2 to alginate flow rate ratio): the CaCl2 concentration examined were 1%, 3% and 5%; the CaCl2 to alginate flow rate ratio examined were 1:1, 2:1, 3:1, 4:1 and 5:1.

The results unit was um.Structural stabilityParameter values that achieved strut formability (listed in Table 2.) were further utilized to fabricate the 3D porous laden structure (Figure 7.). Structural stability test were performed on the fabricated scaffolds. The structure before and after fluidic disturbance were showed in Figure 8- 10. From the observation, the fabricated scaffold would be able to hold its structure in both 4.5% and 6% alginate concentration that examined (Figure 8.). In material composition examination, the scaffold would be able hold the fluidic disturbance in the material composition of pure A0682 and 3:1 mixture (A0682 to 71238). The scaffolds fell apart in the material composition of 2:1, 1:1 mixture (A0682 to 71238) and pure 71238 (Figure 9.). When cross-linking process parameters were examined, all the scaffolds succeeded to sustain the scaffold structure against fluidic disturbance in the cross-linking range examined (Figure 10.).Side View Figure 7.

3D porous laden structure fabricated with in situ printing method using default fabrication parameters.Top ViewAngled ViewStructural stability in regarding to alginate concentration (viscosity) 4.5% Alginate6% AlginateBefore Disturbance After Disturbance Figure 8. Structural stability in regarding to alginate concentration: the alginate concentration examined were 4.5% and 6% (w/w). The structures were stable in both alginate concentrations.After DisturbanceBefore Disturbance1645Structural stability in regarding to material compositionPure A06823:1 (A0682: 71238)Before DisturbanceAfter DisturbanceBefore DisturbanceAfter Disturbance2:1 (A0682: 71238)1:1 (A0682: 71238)Before DisturbanceAfter DisturbanceBefore DisturbanceAfter DisturbancePure 71238Before Disturbance Figure 9. Structural stability in regarding to material composition: the material compositions examined were Pure A0682; 3:1; 2:1; 1:1 (A0682: 71238) and pure 71238.

Structural stability in regarding to material composition: the material compositions examined were Pure A0682; 3:1; 2:1; 1:1 (A0682: 71238) and pure 71238. The structures were stable only in pure A0682 and 3:1 (A0682: 71238) material composition.After disturbanceStructural stability in regarding to CaCl2 to Alginate flow rate ratio in 3% CaCl23% CaCl2 with 3X flow rate ratio3% CaCl2 with 4X flow rate ratioBefore DisturbanceAfter DisturbanceBefore DisturbanceAfter Disturbance3% CaCl2 with 5X flow rate ratio1646Before DisturbanceAfter DisturbanceStructural stability in regarding to CaCl2 to Alginate flow rate ratio in 5% CaCl25% CaCl2 with 2X flow rate ratio5% CaCl2 with 3X flow rate ratioBefore DisturbanceAfter DisturbanceBefore DisturbanceAfter Disturbance5% CaCl2 with 4X flow rate ratio5% CaCl2 with 5X flow rate ratioBefore Disturbance After Disturbance Figure 10.

Structural stability in regarding to cross-linker availability (CaCl2 concentration and CaCl2 to alginate flow rate ratio): the examined CaCl2 concentrations were 3% and 5%; the CaCl2 to alginate flow rate ratios were 2:1; 3:1; 4:1 and 5:1. The structures were stable in all the cross-linker parameters.After DisturbanceBefore DisturbanceStructural integrityThe fabricating parameters (Table 3.) that succeeded to form the strut and sustain the structure stability were further carried out for structural integrity. The scaffolds that incubated in cell culture medium were checked every two days from day 1 to day 15. From the observation, the fabricated scaffolds kept its structure when 3:1 (A0682: 71238) material composition was selected. The scaffolds that fabricated by pure A0682 was degraded out at day 3 (Figure 11.).

The scaffolds that fabricated by pure A0682 was degraded out at day 3 (Figure 11.). There is no noticeable structure disturbance for other fabricating parameters when 3:1 (A0682: 71238) material composition was utilized.Structural integrity in regarding to material composition Pure A06821647Day 0Day 1Day 3 3: 1 (A0682: 71238)Day 0Day 1Day 3Day 5Day 7Day 13 Figure 11. Structural integrity in regarding to material composition: The fabricated scaffolds were incubated in cell culture medium and checked from day 0 to day 15 every two days. Scaffold can hold its porous structure for 15 days when 3:1 (A0682: 71238) material composition was selected but degraded out at day 3 when pure A0682 was used.Day 9Day 11Day 15DiscussionIn the solid freeform fabrication-based in situ printing process, struts are first fabricated as basic element. A pattern of the struts were designed on each layer and were stacked together to form the porous 3D scaffold.

A pattern of the struts were designed on each layer and were stacked together to form the porous 3D scaffold. To produce a porous scaffold with controllable structure feature and decent stability and integrity in cell culture environment, process parameter regarding to three structure characteristics – strut formability, structural stability and structural integrity, were examined.Struts are the basic elements to build the 3D structures. The size of it determines the characteristics of the 3D porous laden structure fabricated (pore size, bio-component accessibility, degradation rate). The strut formation process involved the alginate deposing process and extruded alginate crosslinking process. From the examination of the deposing process parameters (deposition flow rate, nozzle size and travel speed), the strut formability agrees with the developed analytic model: D = ƒ(𝑄, 𝑣) = √4𝑄 𝜋𝑣.

From the examination of the deposing process parameters (deposition flow rate, nozzle size and travel speed), the strut formability agrees with the developed analytic model: D = ƒ(𝑄, 𝑣) = √4𝑄 𝜋𝑣. This result suggested that acylindrical strut can be formed and limited swollen occurred in the fabrication process when cross-linkers are sufficient. When alternating one of the parameters, the strut will not form the expecting size. The cylindrical strut size agreed to the analytic model then is described as strut1648formability achieved. In the examination of the crosslinking process parameters (alginate concentration, crosslinker concentration and crosslinker to deposition flow rate ratio), there is a critical minimum level for alginate concentration and crosslinker availability (crosslinker concentration and crosslinker to deposition flow rate ratio) to ensure the expected strut formation.

The alginate concentration determines both the viscosity of the solution and crosslinking available alginate molecules concentration. The viscosity of the solution governs the spreading behavior of extruded alginate during the crosslinking process, and the crosslinking available alginate molecules concentration determines the stiffness of the strut. When alginate concentration is low, the extruded alginates are prone to spread during cross-linking and less stiff when cross-linked, thus failed to form the expected strut. In the consideration of cross-linker, the CaCl2 concentration and flow rate ratio are dependent parameters both contribute to the cross- linker availability. Strut formability cannot be achieved using 1% Calcium Chloride or 1:1 flow rate ratio (cross-linker to alginate). This result suggested that when one of the parameter is too low, the cross-linker will not be sufficient regardless the other parameter.

This result suggested that when one of the parameter is too low, the cross-linker will not be sufficient regardless the other parameter. It also can be summarized from the results that there is a critical minimum combination of the two parameters. As the two parameters lied higher than the critical combination, strut formability can be achieved (Figure 6D). Thus, in the strut fabrication process, as the minimum crosslinking process sufficed, a strut can be fabricated based on a developed analytic model. Further, it can be observed that the printing system delivered a precise strut size within 10 um as formability achieved.When examining cell behavior in vitro, it is essential that the in vitro model can stand decent fluidic disturbance when different biological assay were applied or fluidic flow that mimic the circulation system in vivo were applied.

Similarly, cell reaction to stimulus is a time depending process, therefore a scaffold product that can sustain it structure within a decent time frame is necessary. To render a scaffold product with decent stability and integrity in cell culture environment, the structural stability and integrity studies were performed. From the results of the two studies, the stability and integrity was only affected by the alginate composition. Two different types of alginate were selected and mixed in different ratios as the parameter of alginate composition. Scaffold fabricated using alginate with low viscosity and low glucuronic acid content (A0682) was found more stable as structure against fluidic disturbance but degraded quickly in cell culture environment. The product with medium viscosity and high glucuronic acid content (71238) can hold its structure longer in cell culture environment but less stable against gentle fluidic disturbance.

The product with medium viscosity and high glucuronic acid content (71238) can hold its structure longer in cell culture environment but less stable against gentle fluidic disturbance. The mixture ratio of 1:3 (71238 to A0682) delivered the scaffold product with decent stability and integrity.1649Reference[1][2]R. Langer and J. P. Vacanti, "Tissue Engineering," Science, vol. 260, pp. 920-926, May 14 1993. Y. L. Cao, J. P. Vacanti, K. T. Paige, J. Upton, and C. A. Vacanti, "Transplantation of chondrocytes utilizing a polymer-cell construct to produce tissue-engineered cartilage in the shape of a human ear," Plastic and Reconstructive Surgery, vol. 100, pp. 297-302, Aug 1997.[3] W. C. Puelacher, J. Wisser, C. A. Vacanti, N. F. Ferraro, D. Jaramillo, and J. P. Vacanti,[4][5][6][7][8]"Temporomandibular-Joint Disc Replacement Made by Tissue-Engineered Growth of Cartilage," Journal of Oral and Maxillofacial Surgery, vol. 52, pp. 1172-1177, Nov 1994. Y. L. Weng, Y. L.

 Vitro Assessment of Laser Sintered Bioactive Glass Scaffolds with Different Pore GeometriesKrishna C. R. Kolan1, Albin Thomas1, Ming C. Leu1, and Gregory E. Hilmas21Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409 2Department of Materials Science and Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractThe pore geometry of bioactive glass scaffolds intended for use in bone repair or replacement is one of the most important parameters that could determine the rate of bone regeneration. The pore geometry would also affect the mechanical properties of the scaffolds and their rate of degradation. Scaffolds with five different architectures, having ~50% porosity, were fabricated with silicate (13–93) and borate (13–93B3) based bioactive glasses using a laser sintering process.

Scaffolds with five different architectures, having ~50% porosity, were fabricated with silicate (13–93) and borate (13–93B3) based bioactive glasses using a laser sintering process. An established, late-osteoblasts/early-osteocytes cell line was used to perform cell proliferation tests on the scaffolds. The results indicated that the cells proliferate significantly more on the scaffolds which mimic the trabecular bone architecture compared to traditional lattice structures.1. IntroductionThe discovery of Bioglass® by Prof. Hench in 1969 has led to active research interest in the field of bioactive materials in the past four decades [1, 2]. The Bioglass® is a commercially available bioactive glass containing mainly SiO2, Na2O, and CaO. Over the years, other beneficial elements were added to the base ternary system and the weight percentage of the elements has changed to develop new glasses which offer excellent bioactive characteristics compared to glass-ceramics or ceramics [3].

13-93 glass is a silicate based bioactive glass which is FDA approved and has a relatively high SiO2 content compared to 45S5 glass. Although the rate of degradation slows down due to the higher SiO2 content, 13-93 glass has been shown to have superior mechanical properties, promote growth and differentiation of osteoblasts in vitro and bone growth in vivo for repair of a critical size bone [4-6]. The molar concentration of SiO2 in 13-93 glass has also been replaced by B2O3 in the 13-93B3 glass formulation. Borate based glasses are chemically less durable than silicate based glasses, and therefore degrade at a faster rate and allows faster bone formation [7]. Borate based bioactive glasses not only bond to the surrounding hard tissue but are also known to bond with soft tissues [7]. Table 1 shows the compositions of the borate based 13-93B3 bioactive glass versus the silicate based 13-93 bioactive glass. The original Bioglass® formulation by Prof.

Table 1 shows the compositions of the borate based 13-93B3 bioactive glass versus the silicate based 13-93 bioactive glass. The original Bioglass® formulation by Prof. Hench, termed 45S5 glass, is also presented for comparison.Table 1. Compositions (in wt.%) of 13-93 and 13-93B3 glasses compared to 45S5 glass.45S5 13-93 13-93B3SiO2 45 53 -P2O5 6 4 3.7CaO 24.5 20 18.5MgO - 5 4.6Na2O 24.5 6 5.5K2O - 12 11.1B2O3 - - 56.6Porosity, pore size, and pore geometry are some of the important aspects of a scaffold for bone repair and they affect material degradation and mechanical properties. Although porosity and pore size can be controlled using the extrusion based additive manufacturing (AM) processes like the Freeze-form Extrusion Fabrication [8] and Robocasting [9-11] processes, these processes provide very limited control over the pore geometry in fabricating porous parts because of the layer-by-layer filament deposition in the fabrication process.

In comparison, powder-bed based AM techniques like the selective laser sintering (SLS) or selective laser melting (SLM) processes provide flexibility in fabricating porous parts with complex pore geometries and lattice structures as they do not require support structures during part fabrication, which allows for better control of part geometry at a microstructural level (10’s of microns). In our previous work, we have shown that 13-93 glass scaffolds fabricated by the SLS process provide good mechanical properties and preferable surface morphology for cell proliferation [12, 13]. Our investigation on the effects of pore geometry on the mechanical properties of the scaffolds, and their degradation behavior when immersed in simulated body fluid, suggested that pore geometry significantly affects the scaffold’s mechanical properties at lower porosities compared to at higher porosities [14].

The technical literature contains only limited studies on the effect of pore geometry on cell proliferation. Some of the studies indicate that the cells react to the local curvature of the pore and the permeability of the implant in vivo [15, 16]. Additionally, more recent studies have indicated that complex pore geometry for the scaffold aids in achieving good cell seeding efficiency [17, 18]. Overall, it is still unclear to the scientific community as to how pore geometry affects tissue growth. Therefore, AM processes provide an opportunity to investigate the effects of scaffold pore geometry in vitro.In the current work, we investigated the effects of pore geometry in the bioactive glass scaffolds on cell proliferation in a static cell culture. Scaffolds made with 13-93 glass and 13- 93B3 glasses were fabricated with ~50% porosity using five different pore geometries.

Scaffolds made with 13-93 glass and 13- 93B3 glasses were fabricated with ~50% porosity using five different pore geometries. The fabricated scaffolds were seeded with bone cells and incubated at ~37°C for 2, 4, and 6 days followed by MTT assays to quantify the metabolically active cells. Section 2 describes the materials and methods used for this study. Section 3 presents and discusses the cell proliferation results obtained for the 13-93 and 13-93B3 scaffolds.2. Materials and Methods2.1. Fabrication of scaffolds13-93 and 13-93B3 bioactive glasses (prepared by Mo-Sci Corp., Rolla, MO) with an average particle size of 12 µm (d50 – 50% of the particles are less than 12 μm) and d90 of 30 µm were used in this research. Particle size distributions were obtained using a laser diffraction- based particle size analyzer (S3500, Microtrac Inc., Largo, FL).

Particle size distributions were obtained using a laser diffraction- based particle size analyzer (S3500, Microtrac Inc., Largo, FL). The glass particles were mixed with stearic acid as the sole binder (C18H36O2, grade HS, Acros Organics, Morris Plains, NJ) and dry ball-milled for 8 hrs with ZrO2 milling media to obtain the feedstock powder for the SLS machine. A binder content of 15% by weight was used for this purpose as this quantity provides the best possible results and is based on our previous study of SLS using 13-93 bioactive glass [19]. The fabrication experiments were carried out on a commercial DTM Sinterstation 2000 machine. The effect of SLS parameters on fabricating scaffolds using stearic acid binder and bioglass powder has been investigated in our previous work. The same set of parameters (laserpower – 5 W, scan speed – 508 mm/s, scan spacing – 0.23 mm, layer thickness – 0.076 mm) were adopted for the current study [19].

The same set of parameters (laserpower – 5 W, scan speed – 508 mm/s, scan spacing – 0.23 mm, layer thickness – 0.076 mm) were adopted for the current study [19]. The fabricated green parts were post-processed in a three-stage programmable air furnace (Vulcan Benchtop Furnace, York, PA). The following heat treatment schedule was used for this study: de-binding at a heating rate of 0.1°C/min to 550°C; heating rate increased to 1°C/min until a final sintering temperature of 570°C with a 1 hour hold for 13-93B3 scaffolds (700°C and 1 hour hold for 13-93 scaffolds) and then the furnace was turned off and allowed to cool naturally to room temperature. Optical microscopy was used to measure the pore sizes and dimensions of the sintered scaffolds. Scaffolds measuring 1 cm x 1 cm and ~3 mm thick containing at least two rows of the lattice unit were used in the experiments so that the effect of pore geometry could be investigated.

Scaffolds measuring 1 cm x 1 cm and ~3 mm thick containing at least two rows of the lattice unit were used in the experiments so that the effect of pore geometry could be investigated. The scaffolds were ultrasonically cleaned in distilled water 3 times for 5 minutes each, followed by a similar cleaning procedure using ethyl alcohol. 13-93B3 scaffolds were only cleaned with ethyl alcohol as they are known to be highly reactive with water. The scaffolds were then dry-heat sterilized overnight at ~200°C before using them for cell proliferation tests.2.2. Cell proliferation testsThe established MLO-A5 line of mouse late-osteoblast, early-osteocyte cells was used as a model for the in vitro tests to determine cell growth [20]. The cells were cultured in phenol red free α-MEM medium supplemented with 5% fetal bovine serum and 5% new born calf serum (Life Technologies, Grand Island, NY).

The cells were cultured in phenol red free α-MEM medium supplemented with 5% fetal bovine serum and 5% new born calf serum (Life Technologies, Grand Island, NY). Scaffolds were placed on a sterilized teflon sheet and seeded with 80,000 MLO-A5 cells suspended in 40 μl of complete medium. After a 2 h incubation to allow cell attachment, the scaffolds with attached cells were transferred to a 12- well plate with 2 ml of complete medium per well and incubated at 37°C in a humidified atmosphere of 5% CO2. The media was replaced every two days during the incubation period. The cell-seeded scaffolds were placed in a medium containing the tetrazolium salt MTT (50 μg per 250 µl of medium) for the last 4 h of incubation to permit visualization of metabolically active cells. The samples were rinsed with phosphate buffered saline (PBS) at the conclusion of the incubation period and blotted dry.

The samples were rinsed with phosphate buffered saline (PBS) at the conclusion of the incubation period and blotted dry. The insoluble purple formazan, which is the product of mitochondrial MTT metabolism, was then extracted from the scaffolds with 2.0 ml of ethyl alcohol and measured at a wavelength of 570 nm using a spectrophotometer (Evolution 300 UV- Vis, Thermo Scientific, Bannockburn, IL).2.3. Statistical analysisAt least three samples in each set (2, 4, and 6 days) were used for cell proliferation tests, and the absorbance data collected from the spectrophotometer is reported as mean ± SD. Analysis for differences in absorbance after the incubation period and among different architectures was performed using one-way ANOVA with Tukey’s post hoc test. Differences were considered significant for P < 0.05.3. Results and Discussion3.1.

Differences were considered significant for P < 0.05.3. Results and Discussion3.1. Design and fabrication of 13–93 and 13–93B3 scaffoldsThe CAD models of the repeatable units of the five architectures considered in this work are shown in Figure 1(a). The first unit cube, termed as a “cubic” architecture, is modeled using solid tubular structures which are perpendicular to each other and running in the x, y, and z directions. The second unit cube is formed by subtracting a sphere from a solid cube and is termed as a “spherical” pore architecture based on the geometry of the pore formed in the design. The third unit cube is formed by diagonally joining the corners of a unit cube with tubular structures and is termed the “X” architecture. These three architectures could be termed as typical and frequently used designs in most of the AM techniques to manufacture scaffolds.

These three architectures could be termed as typical and frequently used designs in most of the AM techniques to manufacture scaffolds. The fourth and fifth are “diamond” and “gyroid” architectures, which are based on the freeform surface and try to mimic the trabecular architecture of bone. More about the scaffold architectures, fabrication, and their mechanical properties can be found in our previous work [14]. The fabricated scaffolds, after binder burnout and sintering, are shown in Figure 1(b) as representative specimens from each of the five architectures in the same order as shown in Figure 1(a). All the scaffolds shown in Figure 1(b) have a measured porosity of ~50%. The scaffolds were cut to about ~3 mm thick such that at least two rows of repeatable unit cells would be available to investigate the cell proliferation using the in vitro assessment.Figure 1.

The scaffolds were cut to about ~3 mm thick such that at least two rows of repeatable unit cells would be available to investigate the cell proliferation using the in vitro assessment.Figure 1. (a) CAD models of the repeatable units of five architectures: (i) cubic, (ii) spherical, (iii) X, (iv) diamond, and (v) gyroid; (b) 13-93B3 scaffolds with the five different architectures after sintering.3.2. Cell proliferation on 13–93 scaffoldsFigure 2(a) shows a schematic of the static cell seeding on the scaffolds and how they are kept in a 12-well cell culture plate. At the time of cell seeding, all of the cell suspension volume applied to the scaffolds adhered to the samples since no suspension was observed on the teflon sheet after the scaffolds were aseptically transferred to the 12-well plate.

Optical images of cell- seeded scaffolds with the five different architectures, and incubated with MTT during the last 4 hours of incubation for intervals of 4 and 6 days, are shown in Figure 2(b). The relative intensity of purple formazan staining on these scaffolds was found to increase with the duration of incubation. This is an indication of metabolically active cells undergoing vigorous growth on the scaffolds. The higher cell density (based on the purple color intensity) on the scaffolds with gyroid and diamond architectures in comparison to the scaffolds with cubic, spherical and Xarchitectures can be observed from the optical images of the scaffolds after 4 days of incubation. The scaffolds with the diamond architecture were completely covered with metabolically active cells after 6 days of incubation.

The scaffolds with the diamond architecture were completely covered with metabolically active cells after 6 days of incubation. To quantify the results, the measured absorbance of the amount of formazan product of mitochondrial metabolism that was recovered from the cell-seeded scaffolds labeled with MTT is plotted in Figure 3.Figure 2. (a) Schematic of the static cell seeding on the scaffold with at least two rows of repeating units. The cell seeded scaffolds were later transferred to a 12-well cell culture plate, (b) Optical images showing the MTT labeling of MLO-A5 cells on the on the scaffolds (L-R: cubic, spherical, X, gyroid and diamond) after cell culture intervals of 4 days (top row) and 6 days (bottom row).Figure 3. The absorbance values representing the measurement of cell growth plotted for all the scaffolds after 2, 4, and 6 days of incubation.

The absorbance values representing the measurement of cell growth plotted for all the scaffolds after 2, 4, and 6 days of incubation. Diamond and gyroid architectures show better overall cell proliferation compared to the other architectures.The absorbance values plotted in Figure 3 for different scaffolds after 2, 4, and 6 days of incubation indicate the number of metabolically active cells on the scaffold, thereby directly corresponding to the amount of cell proliferation. To understand the cell proliferation trends from 2 to 4 days, it can be observed from Figure 3 that all the scaffolds provide significant cell proliferation except the scaffold with the X architecture. Further, the diamond and gyroid scaffolds show higher significance (p < 0.005) in comparison to the cubic and spherical scaffolds.

Further, the diamond and gyroid scaffolds show higher significance (p < 0.005) in comparison to the cubic and spherical scaffolds. Whereas the cell proliferation from 4 to 6 days was significant for the gyroid and diamond architecture scaffolds (p < 0.01), the increase in cell proliferation for the other scaffolds was not significant. These results indicate that although the cubic and spherical scaffolds providesignificant growth up to 4 days, the gyroid and diamond scaffolds provide significant growth up to 6 days of incubation, and the scaffold designed with the X architecture provides the least amount of cell growth. To interpret the results among the scaffolds with different architectures based on their incubation period, no significant difference was noticed after 2 days of incubation as the scaffolds have similar values of absorbance. However, the differences were observed after 4 and 6 days of incubation.

However, the differences were observed after 4 and 6 days of incubation. After 4 days of incubation, the scaffolds having the diamond and gyroid architectures offered significantly higher cell proliferation (p < 0.05) in comparison to the X scaffolds. However, this difference was not significant in comparison to the cubic and spherical scaffolds. After 6 days of incubation, the cell proliferation on the diamond and gyroid scaffolds further improved in significance in comparison to the X scaffolds (from p < 0.05 after 4 days to p < 0.005 after 6 days). Also, the cell proliferation on the diamond scaffolds was significantly higher than the cubic and spherical scaffolds (with no significant difference among them after 4 days, changing to significant with p < 0.05 after 6 days).3.2.

Cell proliferation on 13–93B3 scaffoldsThe average absorbance value of the formazan extracted after 2 days of incubation from the 13-93B3 scaffolds was ~0.3 without any significant difference among the scaffolds with different architectures. In comparison, the absorbance values representing the cell growth on 13- 93 scaffolds after 2 days was ~0.4 (Figure 3). The cell proliferation results on the 13-93B3 scaffolds indicate that the amount of formazan recovered from the scaffolds after 4 and 6 days of incubation is decreasing, which is in sharp contrast to the result observed with the 13-93 scaffolds. Figure 4 shows the comparison of the absorbance values of the formazan product recovered from 13-93 and 13-93B3 scaffolds. The absorbance values in almost all the cases after 4 and 6 days were measured to be ~0.2 or less, which is too low for the result to be significant.Figure 4. Comparison of the absorbance values for 13-93 and 13-93B3 scaffolds after 4 and 6 days of incubation.

Comparison of the absorbance values for 13-93 and 13-93B3 scaffolds after 4 and 6 days of incubation. The graphs indicate the reduced number of metabolically active cells on the 13-93B3 scaffolds compared to the 13-93 scaffolds.The results of in vitro assessment of borate based glasses reported in the literature are mixed. There have been some reports on the cytocompatibilty of the borate glass and the osteogenic differentiation of the human mesenchymal stem cells [21]. There have been other studies which reported a reduction in the absorbance value with an increase in the boron ion concentration in the cell culture media [22]. The reduction of metabolically active cells after 4 and 6 days in our study (Figure 4) may be due to the fact that an increase in the boron ionconcentration leads to an increase in the pH value of the culture media (pH – 9). Such an environment could drastically affect the growth of cells.

Such an environment could drastically affect the growth of cells. It should be noted that this result was observed even after replacing the media every day for the 13-93B3 scaffolds in order to reduce the likelihood of a change in pH. In comparison, the media for 13-93 scaffolds was replaced every other day. Though the in vitro assessment of the 13–93B3 scaffolds showed poor cell survival rates, because of the toxicity attributed to boron ion release and its concentration, the results could be regarded as a false negative based on the in vivo assessment of the borate glasses which has been shown to demonstrate markedly faster bone growth and healing, as reported in the literature [7]. In fact, borate glasses have also been shown to even promote angiogenesis and healing in difficult-to-treat soft tissue wounds in diabetic patients [7]. The investigation of the effect of pore geometries in bone growth using the 13–93B3 scaffolds in vivo will be a subject of our future work.3.3.

The investigation of the effect of pore geometries in bone growth using the 13–93B3 scaffolds in vivo will be a subject of our future work.3.3. Effect of pore geometry on cell proliferationApart from the differences in the cell proliferation based on the material used to fabricate the scaffolds, the surface roughness of the scaffold, which is a characteristic of the SLS process and is irrespective of the geometry, could also affect the way cells proliferate on the scaffolds. Our in vitro assessment results indicated that during the initial 2 days of incubation, the difference in the amount of cell proliferation among the scaffolds with different pore geometries did not exist and became significant only after 4 and 6 days. This could be due to the rough surface of all of the scaffolds fabricated by the SLS process. Figure 5a shows a scanning electron microscopy (SEM) image of a strut with rough edges due to the layer-by-layer fabrication process.

Figure 5a shows a scanning electron microscopy (SEM) image of a strut with rough edges due to the layer-by-layer fabrication process. Another SEM image of the surface area surrounding a pore on the scaffold, irrespective of the material and scaffold pore geometry, is shown in Figure 5b. This amount of surface roughness could provide a favorable attachment surface for seeded cells on SLS-based scaffolds. In fact, at the time of cell seeding, all of the cell suspension volume applied to the scaffolds adhered to the samples and no suspension was observed on the teflon sheet after the scaffolds were aseptically transferred to the 12-well plate (100% seeding efficiency). Therefore, during the initial days of incubation, the cells would proliferate on the rough surface of the scaffold as shown in the schematic in Figure 5c.

Therefore, during the initial days of incubation, the cells would proliferate on the rough surface of the scaffold as shown in the schematic in Figure 5c. This could be the reason for not having any significant differences among the different architectures after seeding all the scaffolds with equal number of cells and 100% seeding efficiency. The effects of pore geometry would start to prevail only after the rough surface is covered by the cells, i.e., after 4 and 6 days.Figure 5. (a) The surface of the strut exhibiting layered fabrication in the SLS process, (b) a typical surface of the SLS scaffold, (c) schematic showing the proliferation of cells on the rough surface during the initial incubation period.The effect of pore geometry on tissue growth and the biochemical signals between cells, as well as how the cells react to the radius of curvature of the scaffold, is still a subject of investigation to the scientific community [15, 23, 24].

While some studies have reported that tissue growth relates to curvature, other studies found no effect of curvature on tissue growth; and most of the studies are speculative in nature. In order to comprehend the differences in cell proliferation among the scaffolds in this study, the internal surface area per unit volume was calculated for the scaffolds with different pore geometries, based on their CAD files, since a higher surface area translates to more area for cells to proliferate on the surface. The value for the X architecture (0.7 mm-1) was much lower in comparison to the other architectures (1.4 for cubic, 1.9 for gyroid, and 1.5 for spherical and diamond), and it reflects in the cell proliferation results as the X scaffold offered no significant cell growth (Figure 3). However, the internal surface area per volume alone does not explain the differences in the cell proliferation among the scaffolds.

However, the internal surface area per volume alone does not explain the differences in the cell proliferation among the scaffolds. Though the scaffolds with spherical and diamond architecture have the same value of internal surface area per volume, the cell proliferation results are in sharp contrast in that the cells seeded on the scaffolds with the diamond architecture proliferate significantly more in comparison to the spherical scaffolds. Therefore, the perimeter of a unit cell is plotted with respect to the thickness (height) of the unit cell to understand the changes in the surface area of a unit cell rather than the overall value. Figure 6 shows the variation in the perimeter of the cross section of a unit cell along the thickness, for all the five architectures.

Figure 6 shows the variation in the perimeter of the cross section of a unit cell along the thickness, for all the five architectures. As the length of unit cell is not uniform among the different architectures, the z height is divided by the unit cell length to provide the z value, which varies from 0 to 1 corresponding to the start and finish of the unit cell thickness. Similarly, the perimeter values are also converted to ratios based on the maximum perimeter value for the corresponding unit cell. The curvature of the internal surface could be understood from the variation of the perimeter of the unit cell for different architectures. Among all the scaffolds, the perimeter profile of the diamond architecture resembles a trigonometrical function (sine/cosine) with more frequent changes in sign (positive to negative and vice versa) to the slope of the curve whereas the changes to the slope of the curve are not drastic for the spherical architecture.

Such frequent and drastic changes to the slope of the perimeter profile suggest a larger curvature to the unit cell, which offers a higher cell proliferation as observed from our results (Figure 3). A relatively slow change in slope or a constant slope of the perimeter profile would suggest small or no curvature which appears to be less beneficial for cell proliferation, especially in a static cell culture condition where the nutrient flow is limited and the only external stimulus for the cells to proliferate could be the curvature. In the case of diamond and gyroid architectures, a larger surface area, together with a larger surface curvature, might allow for better cell proliferation in the static cell culture conditions.

In the case of diamond and gyroid architectures, a larger surface area, together with a larger surface curvature, might allow for better cell proliferation in the static cell culture conditions. Therefore, it can be stated that a preferable porous scaffold for tissue engineering applications should provide the necessary surface roughness for the initial cell attachment, and a larger surface area and surface curvature for improved cell proliferation.1038Figure 6. The variation of the perimeter through the length of the unit cell.4. ConclusionsThe in vitro assessment in this study indicates that bioactive glass scaffolds fabricated by the selective laser sintering process provides the necessary surface roughness for the initial phase of incubation for cell attachment.

The cell proliferation study shows the effectiveness of SLS produced, gyroid and diamond architectures in providing sustained cell proliferation for 6 days of incubation compared to other scaffold architectures (cubic, spherical, and X architectures) which did not provide significant cell growth after 4 days of incubation. Our experimental results and analysis indicate that a larger surface area per unit volume of the scaffold, combined with a larger surface curvature, allows for improved cell proliferation in vitro for the scaffold architecture.References[1] Hench, L.L., Wilson, J., An Introduction to Bioceramics (Singapore: World ScientificPublishing Co. Pte. Ltd.) p 47 (1993).[2] Hench, L.L., The story of bioglass. J Mater Sci: Mater Med. 17, 967–978 (2006). [3] Rahaman, M.N., Day, D.E., Bal, B.S., Fu, Q., Jung, S.B., Bonewald, L.F., BioactiveGlass in Tissue Engineering. Acta Biomater.

9, 7025-7034 (2013). Fu, Q., Saiz, E., Tomsia, A.P., Direct ink writing of highly porous and strong glass scaffolds for load-bearing bone defects repair and regeneration. Acta Biomater. 7, 3547- 3554 (2011).[10][11] Michna, S., Wu, W., Lewis, J., Concentrated hydroxyapatite inks for direct-writeassembly of 3-D periodic scaffolds. Biomaterials. 26, 5632-5639 (2005).[12] Kolan, K.C.R., Leu, M.C., Hilmas, G.E., Brown, R.F., Velez, M., Fabrication of 13-93 bioactive glass scaffolds for bone tissue engineering using indirect selective laser sintering. Biofabrication. 3, 025004 (2011).[13] Kolan, K.C.R., Leu, M.C., Hilmas, G.E., Velez, M., Selective laser sintering of 13-93 bioactive glass. Proc. of the 21st Annual Int. Solid Freeform Fabrication Symp. Austin, TX 504-512 (2010).[14] Kolan, K.C.R., Leu, M.C., Hilmas, G., Comte, T., Effect of architecture and porosity on mechanical properties of borate glass scaffolds made by selective laser sintering. Proc.

Acta Biomater. 7 1009-18 (2011).[17][18][19] Kolan, K.C.R., Leu, M.C., Hilmas, G.E., Velez, M., Effect of material, process parameters, and simulated body fluids on mechanical properties of 13-93 bioactive glass porous constructs made by selective laser sintering. Journal of the Mechanical Behavior of Biomedical Materials. 13, 14-24 (2012).[20] Kato, T., Boskey, A., Spevak, L., Dallas, M., Hori, M., Bonewald, L., Establishment of an osteoid preosteocyte-like cell MLO-A5 that spontaneously mineralizes in culture J. Bone Min. Res. 16 1622-33 (2001).[21] Marion, N.W., Liang, W., Liang, W., Reilly, G.C., Day, D.E., Rahaman, M.N., Mao, J.J., Borate glass supports the in vitro osteogenic differentiation of human mesenchymal stem cells. Mechanics of advanced materials and structures, 12(3) (2005).[22]Fu, H., Fu, Q., Zhou, N., Huang, W., Rahaman, M.N., Wang, D., Liu, X., In vitro evaluation of borate-based bioactive glass scaffolds prepared by a polymer foam replication method.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIN-PLANE PURE SHEAR DEFORMATION OF CELLULAR MATERIALS WITH NOVELGRIP DESIGNK. M. Conwaya, S. S. Kulkarnia, B.A. Smitha G. J. Patakya, G. M. Mockoa, J. D. Summersaa Department of Mechanical Engineering, Clemson University, Clemson, SC 29634-0921, United StatesAbstractCellular materials are popular due to their high specific strength, but their in-plane shear behavior is not well understood. Current experimental methods are limited due to the lack of pure shear loading as common arcan-style grips have not been adjusted for cellular materials. A significant concern is a mixture of shear loading with grip induced tension. While in bulk materials the tensile force can be assumed negligible, it has a significant impact on the deformation behavior of cellular materials.

While in bulk materials the tensile force can be assumed negligible, it has a significant impact on the deformation behavior of cellular materials. In this study, finite element modeling simulations were used to demonstrate that using a new sliding grip design reduced grip induced tension on cellular materials. Experimental studies were performed on honeycomb cellular materials with traditional and newly- developed grips to calculate and compare the shear strength and ductility of honeycomb cellular materials.

The study concluded that traditional grips overestimate the shear strength of honeycomb cellular materials and honeycomb cellular materials in pure shear with limited grip induced tension has significantly lower strength and ductility due to the early formation of plastic hinges.IntroductionCellular materials are used in a variety of applications due to their ability to absorb energy [1,2], high specific strength [3,4] and controllable materials properties such as Poisson’s ratio [4– 6], toughness and yield strength [7]. To understand how cellular materials will deform in anticipated loading environments, the mechanical behavior of cellular materials under pure and mixed loads must be understood. The uniaxial tensile and compressive behaviors of cellular materials has been well documented [4,7–10]. Similar to foams, when cellular materials experience low velocity impacts, the material will experience shear loading [11].

Similar to foams, when cellular materials experience low velocity impacts, the material will experience shear loading [11]. However, the pure shear behavior of cellular materials has not been sufficiently studied experimentally, because most grip systems used in shear tests of cellular materials introduce a transverse constraint that adds a tensile component to the shear test [12–17]. Most experimental studies of shear in cellular materials use a three bar double shear test [12,15] or an offset grip single shear test [17], shown in Figure 1. Due to the nature of these grips, the width of the specimen is kept constant. This constraint in the horizontal direction adds a tensile component to the specimen, creating a combined shear and tension test, demonstrated in Figure 2. This combined loading artificially strengthens the specimen as the cellular material is stronger in tension then shear loading.

.~2J--------------Figure 2: The difference in deformation of pure shear and combined shear and tension applied loads. The added tensile component in the combined shear and tension artificially strengthens materials stronger in tension.While the limitations of current cellular material shear deformation grip designs has been discussed in literature [12,17], a grip design has not been adopted that is capable of true pure shear applied loads. Current experimental studies of cellular materials tested in pure shear continue to be combined loading, with the tensile component included [13,15]. Additionally, most numerical studies of shear loading of cellular materials also constrain the height of the specimen as to2297maintain the same loading conditions of experimental studies [14–16,19].

Due to this, the magnitude of the effect the applied tensile load has on the cellular material cannot be quantified through a review of the literature.In this study, traditional offset single shear grips and novel grips that do not constrain the transverse deformation of the cellular material were compared to quantify the effect of induced tension on the shear failure strain of cellular materials. A numerical modeling study of the honeycomb specimens in shear was used to model the differences in specimen behaviors in both loading scenarios and were compared with experimental results.Methods and MaterialsHoneycomb specimens were additively manufactured (AM) by fused deposition modeling (FDM) using AmazonBasic acrylonitrile butadiene styrene (ABS) white 1.75 kg spool filament in a Makerbot Replicator 2X. AM specimens were printed using a 0.4 mm nozzle, a nozzle temperature of 250 °C, a bed temperature of 110 °C, a layer height of 0.2 mm, and a print speed of 90 mm/s.

AM specimens were printed using a 0.4 mm nozzle, a nozzle temperature of 250 °C, a bed temperature of 110 °C, a layer height of 0.2 mm, and a print speed of 90 mm/s. The honeycomb shear specimens were ten cells tall and five complete cells wide with a wall thickness of 1.2 mm, a cell width of 4.35 mm and a total specimen thickness of 3.75 mm, producing a shear specimen with an effective area of 86.4 mm2 using the method to calculate effect area discussed in [7]. After printing, the honeycomb specimens were polished with P120 grit sandpaper up to P1500 grit sandpaper.Traditional offset single shear grips, referred to as the fixed-fixed grips in the study, were designed for the honeycomb shear specimens that kept the width of the specimen constant throughout testing, as can be seen in Figure 3a, inducing mixed loading on the specimens.

Novel grips, referred to as the sliding-fixed grips (Figure 3b), were designed with a channel that allowed the grips to translate in the transverse direction during the test. A machinist at Clemson University manufactured both grips out of 316L stainless steel. A plastic shim with made of poly-lactic acid with an approximate coefficient of friction of 0.25 [20] was added to the channel of the sliding grips to minimize sliding resistance in the grip.Figure 3: a) Offset single shear fixed-fixed grips. The width of the specimen is maintained throughout the shear test, inducing a tensile load on to the cellular material. b) Novel sliding- fixed grips that allow the grips to translate in during the test to eliminate the applied tensile load seen in the offset grips.2298The grips were loaded in tension, thereby loading the cellular specimen in shear. The cellular shear specimens were tested to failure in two grip combinations: fixed-fixed and sliding- fixed.

The cellular shear specimens were tested to failure in two grip combinations: fixed-fixed and sliding- fixed. Experiments were performed as displacement-controlled (cid:68)(cid:87)(cid:3)(cid:68)(cid:3)(cid:71)(cid:76)(cid:86)(cid:83)(cid:79)(cid:68)(cid:70)(cid:72)(cid:80)(cid:72)(cid:81)(cid:87)(cid:3)(cid:85)(cid:68)(cid:87)(cid:72)(cid:3)(cid:82)(cid:73)(cid:3)(cid:22)(cid:19)(cid:3)(cid:541)(cid:80)/s using an MTS Landmark 370 hydraulic load frame. A dual camera system was used for a multiscale understanding of the deformations. One camera was focused on the deformation of a single cellular unit cell while the other captured the global deformation of the entire specimen. Both cameras were Point Grey model GS3 and captured images at a rate of 1 Hz. Digital image correlation (DIC) was used to calculate the strains in the legs of the specimens during deformation.

Both cameras were Point Grey model GS3 and captured images at a rate of 1 Hz. Digital image correlation (DIC) was used to calculate the strains in the legs of the specimens during deformation. The camera capturing the deformation of a single unit cell was positioned directly in front of the specimen and was equipped with a Navitar lens and a Navitar 1X adaptor with a resolution of 120 pix/mm. The camera capturing the global deformation was positioned next to the first camera, slightly offset from the specimen and was equipped with a Schneider Kreuznach Xenoplan lens model 1001960. Images were used to calculate the full field displacements and strains using the commercial DIC software, VIC 2D. Strain calculations were performed with a virtual strain gage of 101 and a spatial resolution of 121 following the procedure outlined in [21].

Strain calculations were performed with a virtual strain gage of 101 and a spatial resolution of 121 following the procedure outlined in [21]. Before testing, specimens were mechanically polished and speckled using an Iwata Custom Micro airbrush model CM-B2 and opaque black Testors Aztek airbrush paint, 9441A.Numerical modeling simulated specimen loading scenarios using the finite element software ANSYS Workbench 17.0. Grips and specimens were modeled using Solidworks modeling software. The behavior of ABS is simulated using the Bergstrom-Boyce model to predict the non-linear viscoelastic response of the ABS elastomer. Tensile stress-strain data from FDM ABS was used to calibrate the material model. The material properties used in the simulation are summarized in Table 1. The geometries were meshed using the adaptive size function in Workbench, with coarse relevance center, medium smoothing and fast transition.

The geometries were meshed using the adaptive size function in Workbench, with coarse relevance center, medium smoothing and fast transition. The mesh settings result in a total of 6101 and 5925 solid elements (SOLID 186 & SOLID 187) for the fixed and sliding cases respectively. Specifically, the honeycomb structure is meshed using 5265 and 5323 quadratic tetrahedral Solid 187 elements (10 nodes), with three degrees of freedom, for the fixed and sliding cases respectively. The contact between the solid components is meshed using CONTA 174 and TARGE170 element pair. The boundary conditions used for modeling the two grips are summarized in Table 2. The system was modeled in tension, thus loading the cellular specimen in shear using a static structural analysis that did not consider strain rate dependent loading. Force displacement data from the experimental tests has been used to find the displacement value at which failure occurs.

Force displacement data from the experimental tests has been used to find the displacement value at which failure occurs. The distance is averaged over three test values.To model both grips systems, appropriate boundary conditions were applied to mimic the conditions experienced by the physical grips. The fixed grip system was modeled by completely constraining the bottom grip, displacing the top grip in the vertical direction, and completely constraining rotation of the top plane of the grip. The bottom grip of the sliding system was also completely constrained, and the top grip displaced in the vertical direction, without rotation. However, a frictional contact (μ=0.3) was added to the system between the horizontal and vertical components of the grips to allow the grips to freely translate together and apart as the specimen is loaded.

The boundary conditions can be seen in Figure 4.2299Table 1: Material Properties used in simulationProperty Density [kg/m3] Young’s Modulus [GPa] Poisson’s ratio Bulk Modulus [GPa] Shear Modulus [GPa] Yield Strength [MPa] Ultimate tensile strength [MPa]FDM ABS 1050 1.54 0.42 4.17 8.93 29 30316L Stainless Steel 7850 200 0.30 167 76.29 250 460Table 2: Boundary conditions used in simulation for fixed and sliding gripsPart Upper bracket Lower bracket X axis separation ConstantFixed grips Displacement (x=0, y=-6.8 mm, z=0) Displacement (x=0, y=-6.3 mm, z=0) Fixed supportSliding gripsFixed support Frictional sliding with coefficient 0.3Fixed gripsSl iding gripsFixed SupportB DisplacementFixed SupportB DisplacementFixed Support 2Frictional Contact0.000 -0.01S0.030 -0.04S0.060(m)0.000 -0.0 180.03S -0.0S30.070(m)Figure 4: Boundary conditions of fixed and sliding grip systems in ANSYS modelResults and DiscussionSimulation of Both GripsThe deformation of honeycombs specimens in both the fixed and sliding grips were simulated to predict how local strains and forces would change between the grip systems.

Table 3 shows the results for the fixed-fixed grips and the sliding-fixed grips. Comparing the strain values2300between the sliding and fixed grips, the sliding grips exhibit an overall reduction in tensile strain with a maximum reduction of 23.08%. From this reduction in tensile strain in the sliding-fixed grips the simulation suggests that the fixed-fixed grips cause mixed loading on the specimen instead of pure shear and the sliding-fixed grips apply loading closer to pure shear.Experimental Results of Fixed-Fixed GripsHoneycomb specimens were tested to failure in the fixed-fixed grips. The specimens deformed in one of three manners. The first distinct way specimens failed was a uniform deformation of the entire specimen where all of the unit cells of the specimen would transform from hexagons to rotated rectangles as seen in Figure 5b. This uniform deformation is the deformation method captured by the simulation.

This uniform deformation is the deformation method captured by the simulation. During this deformation, plastic hinges formed at the intersections of the cell walls, as indicated by the large local shear strains shown in Figure 6.Figure 5: Honeycomb offset single shear specimen. Specimens failed in three distinct manners. a) Undeformed specimen. b) Uniform deformation: all hexagonal unit cells transform to rectangles with applied global shear loading. c) Deformation of cells at grip: Minor uniform deformation of hexagonal unit cells, horizontal cell wall grip fails, initiating immediate failure in entire column at wall. d) Major deformation in one column of cells in the specimen, other cells in specimen only experience minor deformation. Camera is offset, not square to specimen, to allow the camera capturing the close up unit cell deformation to be square.

Camera is offset, not square to specimen, to allow the camera capturing the close up unit cell deformation to be square. This is causing an optical illusion suggesting the specimen is deforming out of plane; it is not.23010.02100.014Figure 6: Plastic hinges formed at joints, as indicated by the large local shear strains, of the cellular material as the unit cells deformed from hexagons to rotated rectangles during global shear deformationThe second and third deformation modes involved the honeycomb failing at a single column. The specimen failed in a single column either at the wall of the specimen, Figure 5c, or at a column in the center of the specimen, Figure 5d. When the specimen failed at the column at the wall of the specimen, as seen in Figure 5c, the hexagonal unit cells did not deform to the degree seen in the uniform deformation case, Figure 5b. Instead the horizontal walls in the column attached to the specimen wall would develop plastic hinges and fail [18].

Instead the horizontal walls in the column attached to the specimen wall would develop plastic hinges and fail [18]. As often seen in cellular materials, when one wall or unit cell fails, the sudden increased loads on the neighboring cell walls leads to plastic collapse of the cellular material where the surrounding walls fail and the material fractures along a row of cells [9,10,15,18]. Similarly, this deformation method can occur in a column not adjacent to the specimen representing the third deformation method, as in Figure 5b. This occurs when one column deforms to a greater degree than its surrounding columns, so that the originally horizontal walls of the cells in the center column of the specimen rotate as the specimen is deformed, however the neighboring columns experience noticeably less deformation.

The failure in one cell wall of the specimen that leads to the plastic collapse of the entire column is due to a stress concentration that causes the lack of integrity of that cell wall. In AM materials internal voids and printing defects will introduce stress concentrations throughout the build. These introduced stress concentrations during the build process explain the variation in the performance of the cellular specimens. Additionally noticeable crazing is present at the plastic hinges that develop as the cellular specimen deforms, further deforming areas with already pre-existing stress concentrations [22]. The variability in deformation modes between the specimens could be due to variation of defects from printing between the specimens. The surface of each specimen was polished to minimize the effects of surface flaws, however the inner wall of each honeycomb was not polished and surface flaws could have been present at the hinge points in the specimens.

,J,;. ,.,rx' :e 600>,4000.00502000 .0050.01200400800800 1000 1200 1400 1600 1800X (pix)txx" Strains0 .0 2txy" Strains0.0 1100010000 .0088000.018000 .0 06'x :e. 6000.005'x s 6000.004>,>,4004000 .00 22000 .00520000 .0 1200400600 800 1000 1200 1400 1600 1800200400600800 1000 1200 1400 1600 1800X (pix)x(pix)Figure 8: Strain transformations aligning with cell wall 30° CW off vertical of uniformly deformed offset single shear specimen2303Strains calculated through DIC were rotated using the strain-transformation equations in Equations 1-3 to align with the angled cell walls of the honeycomb in the cellular specimen, Figure 9, using a MATLAB program. The rotated strains for the uniformly deformed specimen, Figure 5b, shown in Figure 7 and Figure 8.

As the rotated rectangle is lengthened, the “long edges” of the rectangle are straightening due to the tensile component of the mixed load induced onto the cellular material by the grips. Additionally the cell walls that make up the “long edge” of the rectangle begin to bend to resist this deformation, as indicated by the large shear forces in the “long edge” walls in Figure 7c. As the honeycomb unit cell transitions to the rotated rectangle, a Poisson’s effects of the rectangle narrowing causes the “short edge” of the rotated rectangle above and below to be loaded in tension as it resists this movement as indicated by the positive tensile loads in the “short edge” walls in Figure 8a.

Additionally, plastic hinges form along with visible crazing at the triple junction in the middle of the “long edge” of the rectangle internal to the rectangle as the hinges are plastically deformed, resisting this deformation.2304(1)(2)(3)A comparison of the DIC results and the ANSYS model are shown in Figure 10. The model has good agreement with the DIC results for the magnitude and location of the local shear (cid:2013)(cid:3051)(cid:3052)(cid:1314) strains, however the model disagrees with the DIC data on the location of the local (cid:2013)(cid:3052)(cid:3052)(cid:1314) strains.

The model has good agreement of the local (cid:2749)(cid:1824)(cid:1825)(cid:1314) shear strains and fair agreement on the location of plastic hinges and concentration of (cid:2749)(cid:1825)(cid:1825)(cid:1314) strains2305The local shear strains of the fixed grips specimen show that the specimen is developing plastic hinges where two arms intersect as well as the arms of the specimen are bending. The formation of hinges is the dominate behavior, evidenced by the larger strains at the hinges as well as the eventual failure of the specimen at the hinges. As the specimen is loaded in shear, the transformation of the hexagonal honeycomb shape to rotated rectangles compresses the sides of the hexagon that form the long edge of the rectangle as well as the short end of the rectangle due to Poison’s ratio as shown by the negative (cid:2013)(cid:3052)(cid:3052)(cid:1314) and (cid:2013)(cid:3052)(cid:3052)(cid:1314)(cid:1314) strains.

Comparing the strain values, the ANSYS model predicted higher values of shear strain ((cid:1837)xy) and normal strain ((cid:1837)yy) when compared to the DIC data. This is attributed to the model not accounting for crazing in the ABS at the hinge locations where the specimen experiences the largest strains. Since the samples are additively manufactured, the material is non-homogeneous and anisotropic, while the simulation model assumes isotropic material, and does not account for crazing and plastic behavior. Due to this reason, the hinge effect is more pronounced in the case of DIC data. The simulation does not account for local defects which could result in non-uniform strain distribution in different hexagon wall locations, resulting in some cells with higher values of strains and some with lower.Sliding-Fixed Shear GripsWhen the honeycomb specimens were tested in the novel sliding grips, the specimen failed in a near pure shear deformation behavior, shown in Figure 11.

The tensile component the sliding grips induced on the specimen was only due to sliding friction and was lower than the fixed grips. In the sliding-fixed grips, the center column of honeycomb cells failed in shear, however the surrounding cells did not transform into the rotated rectangles seen in the fixed grip specimens. A comparison of the final deformation appearance of the sliding grips and fixed grips is shown in Figure 12.Figure 11: Shear deformation behavior of honeycomb specimen in new grips. (a) Specimen before loading (b) Specimen at maximum shear load, single row of cells failed in shear2306Figure 12: Comparison of deformation behavior of (a) sliding and (b) fixed grips. The fixed grips with grip-induced tension stretch cells between the grips forcing cells to transform into rotated rectangle geometries.

The fixed grips with grip-induced tension stretch cells between the grips forcing cells to transform into rotated rectangle geometries. Sliding grips do not stretch cells between grips causing the middle row to fail in a pure shear manner while neighboring cells do not transform.The stress-strain response of the honeycomb specimens in the sliding grips are shown in Figure 13. The sliding grip specimens failed at a significantly lower stress and strain then the fixed grip specimens. This is because the honeycomb cellular specimens are stronger in tension then in pure shear. When the specimen deformed in the fixed grips, the grips cause the specimen to stretch between the grips. This causes all of the honeycomb cells to expand and transform into rotated rectangles.

When the specimen deformed in the fixed grips, the grips cause the specimen to stretch between the grips. This causes all of the honeycomb cells to expand and transform into rotated rectangles. In the sliding grips, the honeycombs did not stretch between the grips and therefore not all of the joints in the honeycomb became plastic hinges, instead only the joints of the horizontal beam that sheared formed plastic hinges. The deformation caused very localized straining to occur on those few joints instead of distributing the strain throughout the entire transforming honeycomb as occurred on the fixed grip specimens. Mixed axial and shear loading causes the honeycomb specimen to fail at a yield strength approximately twice as large as pure shear loading and drastically increases the ultimate strain.

Mixed axial and shear loading causes the honeycomb specimen to fail at a yield strength approximately twice as large as pure shear loading and drastically increases the ultimate strain. Previous studies that have reported results as purely shear that were actually mixed shear and tension may have misreported the ultimate strength of the cellular material by half or more.230754roa.. ~ ..__,. (/) (/) ~ u532-Fixed Grip with Induced Tension Sliding Grip Pure Shear0.000.020.040.06 Strain0.080.10Figure 13: Stress-strain response of fixed and sliding grips. The fixed grips caused a combined loading of tension and shear on the honeycomb specimens. The combined loading artificially strengthened the honeycombs, as they are stronger in tension then in shear.

The combined loading artificially strengthened the honeycombs, as they are stronger in tension then in shear. The sliding grip tests illuminated that honeycomb cellular material is not as strong in shear as the fixed grip test results would otherwise indicate.Local strains of the sliding grip honeycomb oriented with the angles axes of the honeycomb are shown in Figure 14. Plastic hinges form at the top left and bottom right of the horizontal beams. The high localization of strain in those few hinges leads to the early failure compared to the fixed grip specimens where the arms of the honeycomb experienced axial strains in addition to plastic hinging, Figure 7 and Figure 8. Deforming the honeycomb arms axially in the fixed grips as well as rotating the arms decreased the deformation experienced directly by the plastic hinges.

Plastic hinges, indicated by high local strains, formed at the top left and bottom right of the horizontal beam, the location where the beam will ultimately fail2309The simulations of the sliding grips showed good agreement with the DIC as far as the location of the plastic hinges in the sliding grips as shown in Figure 15. The simulation however did not agree on the magnitude of the strains of the sliding grips. As with the fixed grips, this disagreement in magnitude is probably due to the simulation not considering crazing of the ABS honeycombs at the hinges.

The finite element model predicts a reduction in induced tension strain and an increase in the shear strain2310when using the sliding grips (Table 4), demonstrating the differences in the loading effects on the specimens of the two grip designs.Table 3.

Due to this, finite element analysis could be used for future studies to determine the effect of mixed axial and shear loading on other topologies and determine which topologies require experimental study to determine the effect of mixed axial and shear loading and which geometries can be determined through simulation.ConclusionTraditional cellular material shear tests add a tensile component to the cellular materials. While previously assumed negligible, this study showed that the added tensile force artificially strengthened honeycomb cellular materials. Honeycomb cellular materials are stronger in combined tension-shear then pure shear. Under combined loading the entire honeycomb cell transformed to a rotated rectangle, distributing the strain throughout all of the honeycomb cells.

Under combined loading the entire honeycomb cell transformed to a rotated rectangle, distributing the strain throughout all of the honeycomb cells. In pure shear loading, the only part of the cell that deformed was plastic hinges on either side of the horizontal arm of the honeycomb and only the cells in a single column developed plastic hinges, not allowing the strain to be distributed the rest of the cells. Honeycomb cellular material in combined loading is about twice as strong as honeycombs in pure shear, showing that previous studies may have over reported the strength of honeycomb cellular materials in shear.AcknowledgementsThe authors would like to acknowledge experimental facilities provided by the Center for Integrated Nanotechnologies (CINT). Sandia National Laboratories is a multimission laboratory managed and operated by National Technology & Engineering Solutions of Sandia, LLC, a wholly owned subsidiary of Honeywell International Inc., for the U.S.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIn-Process Condition Monitoring in Laser Powder Bed Fusion (LPBF)Mohammad Montazeri*, Prahalada Rao*Mechanical and Materials Engineering Department, University of Nebraska-Lincoln, Lincoln, NE 68588The goal of this work is to monitor the laser powder bed fusion (LPBF) process using an array of heterogeneous sensors. This goal is termed as build status monitoring. The overarching aim is to usher a qualify-as-you-build quality assurance paradigm in LPBF whereby incipient build defects are detected through analysis of data acquired from in-process sensors. This will allow opportune preventive action to avert propagation of build defects across multiple layers.

This will allow opportune preventive action to avert propagation of build defects across multiple layers. In pursuit of this goal, a commercial LPBF machine at the National Institute of Standards and Technology (NIST) was integrated with three types of sensors, namely, a photodetector, high-speed video camera, and SWIR thermal camera with the following objective: to develop and apply a spectral graph theoretic approach to monitor the LPBF build status from the data acquired by the three sensors. This objective will lead to early identification of incipient defects that afflict LPBF despite extensive process automation. The proposed approach is illustrated with experimental sensor data acquired during LPBF of a part having a steep overhang feature of ~ 40.5o. Parts with such steep overhang features may exacerbate deleterious consequences such as poor surface finish, porosity, and distortion.

Parts with such steep overhang features may exacerbate deleterious consequences such as poor surface finish, porosity, and distortion. Hence, close monitoring of the signal patterns during scanning of overhang areas is consequential for early detection of build defects. The proposed approach detected differences between overhang and non-overhang build status for different sensors with the statistical fidelity (F-score) of 95% from thermal camera signatures to 79% with the photodetector. In comparison, conventional signal analysis techniques - e.g., neural networks, support vector machines, linear discriminant analysis, etc., are evaluated with F-score in the range of 40% to 60%.

In comparison, conventional signal analysis techniques - e.g., neural networks, support vector machines, linear discriminant analysis, etc., are evaluated with F-score in the range of 40% to 60%. As part of our forthcoming work, this study will be further expanded to include more build defects, e.g., due to material contamination.Key Words: Laser Powder Bed Fusion (LPBF), Heterogeneous Sensing, In-situ Monitoring, Meltpool Thermal Imaging, High-Speed Imaging, Photodetection, Spectral Graph Theory, Manifold Learning, Laplacian Eigenvectors.Corresponding author: P. Rao, email:rao@unl.edu1Introduction1.1 ObjectiveThe goal of this work is to monitor the laser powder bed fusion (LPBF) process using an array of heterogeneous sensors [1]. This goal is termed as build status monitoring.

This goal is termed as build status monitoring. In pursuit of this goal, a commercial LPBF machine (EOS M270) at the National Institute of Standards and Technology (NIST) was integrated with three sensors, namely, a photodetector (spectral response 300 nm to 1200 nm), high-speed video camera (4000 frames/s), and SWIR thermal camera (1350 nm to 16001264nm, calibrated for measuring temperature in the range 500 °C and 1025 °C) with the following objective.Objective: Develop and apply a spectral graph theoretic approach to monitor the build status in LPBF from the data gathered by the three aforementioned sensors. The intent is to detect the onset of deleterious phenomena, such as thermal-induced distortion [2-5]. These phenomena are precursors to part defects, e.g., geometric inaccuracies [5]. In the worst case, these may ultimately result in build failures.

These phenomena are precursors to part defects, e.g., geometric inaccuracies [5]. In the worst case, these may ultimately result in build failures. The proposed approach is extensible to other AM processes.This objective will answer the following open research question in the context of in-situ sensing and monitoring in LPBF.What sensor signatures (e.g., meltpool shape, spatter pattern) are symptomatic of the build status; and what new analytical approaches are required to extract these signatures from a heterogeneous sensor suite?1.2 Scientific Rationale and HypothesesThe following two hypotheses corresponding to the objective will be tested in this work.Hypothesis 1: The scientific rationale is that each type of build defect in LPBF is connected with specific process phenomena. The onset of these phenomena manifests in statistically distinctive sensor signatures [6-8].

The onset of these phenomena manifests in statistically distinctive sensor signatures [6-8]. Hence, by closely tracking the signatures from in-process sensor data the build status in LPBF can be discriminated. Following this reasoning, we posit that given a signal into a weighted captures the process, e.g., photodetector signal, thermal and high-speed video images, convertingplanar network graph 𝒳𝒳 phenomena in a lower dimensional manifold.withvertices, edges, and weights𝒳𝒳𝐺𝐺(𝑉𝑉, 𝐸𝐸, 𝑊𝑊)𝑉𝑉𝐸𝐸𝑊𝑊The central premise is that despite the transformation of the data into a lower dimensional graph space, a large part of the information in the signal is retained, as shown by Belkin and Niyogi [9- 11]; Sandryhaila et al. [12], Hammond et al. [13]; and Shuman et al. [14, 15].We specifically test the hypothesis in Sec.

[12], Hammond et al. [13]; and Shuman et al. [14, 15].We specifically test the hypothesis in Sec. 2 that by representing the sensor data as an weighted and Eigenvalues and undirected graph ( ) of the graph Laplacian as derived sensor signatures, leads to higher classification fidelity in terms of the F-score — which is an aggregate measure of the Type I and Type II statistical error 𝜆𝜆 — compared to popular parametric approaches, such as neural networks and support vector machines which rely on statistical features from the sensor data [16]., and subsequently using the Eigenvectors𝐺𝐺(𝑉𝑉, 𝐸𝐸, 𝑊𝑊)(𝑣𝑣)Hypothesis 2: Using the proposed graph Laplacian approach, the classification fidelity with the photodetector is within 20% (F-score) of that obtained through the high-fidelity thermal and video cameras. We test this hypothesis in Sec.

We test this hypothesis in Sec. 2.3 by comparing the F-score for each of the three sensors in discriminating between the overhang and non-overhang build conditions.1265Remark – Why Monitoring Overhang Versus Non-Overhang Build Status is a Practically Important Problem.Overhang features typically require support structures, and are in contact with more metal powder material from below compared to solid, non-overhang features. Insufficient or poorly designed support structures is one of the reasons for build failures [4]. Printing of overhang features is a challenging proposition due to the following reasons [3].First, referring to Figure 1 (left), as the thermal conductivity of the powder is roughly one third of a solid part, heat tends to accumulate within the overhang area, i.e., the thermal flux through an overhang is restricted [3].

Constriction of heat to a relatively small area leads to severe thermal gradients within the overhang features which ultimately manifests in distorted builds (termed swelling by Sames et al. [5]). If unchecked, distortion will lead to the build making contact with the recoater mechanism. This contact may damage the recoater/roller and the part; jam the recoater; or even peel the part off the build platen.Second, the impeded heat flux in the overhang may lead to overheating which affects the surface quality and microstructure of the downward facing overhang sections [7]. A practical example of this effect in LPBF of a 316L stainless steel knee implant having an overhang section is evident in Figure 1 (right).Third, Creaghs et al. observed during scanning of the overhang the aspect ratio (length to width) of the meltpool increases [8].

observed during scanning of the overhang the aspect ratio (length to width) of the meltpool increases [8]. As a consequence, the authors posit, the meltpool becomes unstable due to onset of a phenomena called Plateau-Rayleigh effect and splits into discrete chunks. This disjointed nature of the meltpool is believed to be at the root of the poor surface finish and porosity observed in overhang sections [3]. These effects for building parts with unsupported overhangs was theoretically confirmed by King et al. [17]Figure 1: (left) Illustration of restricted heat flux in the overhang features. (right) Stainless steel 316L knee implant built on a production LPBF machine showing significant overheating and poor surface finish in overhang areas.12662 Proposed MethodologyApplication of graph theoretic approaches for signal processing is a nascent domain with recent notable review articles by Hammond et al. [13], Sandryhaila et al. [12], and Shuman et al. [14, 15].

[13], Sandryhaila et al. [12], and Shuman et al. [14, 15]. Graph theoretic shift operators and signal transforms are in their infancy [12-15, 18, 19]. Niyogi et al. [9-11] in a series of seminal articles proposed embedding high dimensional data as an undirected graph, and subsequently projecting the data into the Eigenvector space of the graph Laplacian[9-11]. A conference on signal processing on graphs, with implications to Big Data was recently organized [20].The convention is to denote matrices and vectors with bold typesets. Suppose we gather a (time series or images) from a process. Further, consider that the ,.

Suppose we gather a (time series or images) from a process. Further, consider that the , . In LPBF these states could refer to different printing conditions, such as non- 𝓈𝓈2 𝓈𝓈1 associated from which an , we want to find the process state 𝓈𝓈𝑖𝑖 ,sequence of sensor data, process manifests indifferent known process conditions or build states labelled as𝔁𝔁𝑖𝑖 =⎡ ⎢ ⎢ ⎣1 𝑎𝑎1 1 𝑎𝑎𝑡𝑡 ⋮ 1 𝑎𝑎𝑚𝑚2 𝑎𝑎1𝑑𝑑 ⋯ 𝑎𝑎1 𝑑𝑑 𝑎𝑎𝑡𝑡 ⋮ 𝑑𝑑 ⋯ 𝑎𝑎𝑚𝑚⋱⎤ ⎥ ⎥ ⎦The approach involves three phases; these are described herewith.Step 1: Transform the signals weighted network graph 𝔁𝔁𝑖𝑖 W is the weight between the edges. 𝐺𝐺𝑖𝑖(𝑉𝑉, 𝐸𝐸, 𝑊𝑊)into an undirected, are the vertices and edges of the graph and 𝓈𝓈𝑖𝑖corresponding to each pre-labeled stateand. Where,𝐸𝐸𝑉𝑉 is computed from the graph.

𝐺𝐺𝑖𝑖(𝑉𝑉, 𝐸𝐸, 𝑊𝑊)into an undirected, are the vertices and edges of the graph and 𝓈𝓈𝑖𝑖corresponding to each pre-labeled stateand. Where,𝐸𝐸𝑉𝑉 is computed from the graph. The first non-zero re used as an orthogonal basis set corresponding to theStep 2: The spectral graph Laplacian matrixgraph Laplacian Eigenvectors𝓛𝓛𝑖𝑖𝐺𝐺𝑖𝑖.process state 𝓃𝓃 Step 3: Each a set of coefficients𝒗𝒗𝑖𝑖 a is decomposed by taking an inner product called graph Fourier coefficients.𝓈𝓈𝑖𝑖akin to a Fourier transform intoT𝒗𝒗𝑖𝑖𝒙𝒙𝑖𝑖 in block matrix ,𝔁𝔁𝑖𝑖 • The graph Fourier𝒄𝒄𝑖𝑖coefficientsform ,asis written. The ℂ =corresponding to different statesis called the dictionary. [𝒄𝒄2 ] ⋯ [𝒄𝒄𝑛𝑛matrix T [[𝒄𝒄1 ]TTT𝓈𝓈𝑛𝑛𝓈𝓈𝑖𝑖 ⋯] ⋯ [𝒄𝒄𝑖𝑖𝓈𝓈1𝓈𝓈2 is taken with each of the n basis are called the candidate in terms of the]]Step 4: Given an unlabeled signal vector sets one at a time; where n are the different states. The matrices 𝑦𝑦 coefficients.

The matrices 𝑦𝑦 coefficients. Each, an inner productℂT𝓹𝓹𝑖𝑖 = 𝑦𝑦 is compared with the corresponding𝓋𝓋𝑖𝑖 in the dictionary T 𝓹𝓹𝑖𝑖 ℂ𝓋𝓋𝑖𝑖 in the dictionary T 𝓹𝓹𝑖𝑖 ℂT𝒄𝒄𝑖𝑖𝓹𝓹𝑖𝑖1267squared error Experimental Data from the Laser Powder Bed Fusion Process. The comparison resulting in the least error is the designated Application to𝑒𝑒𝑖𝑖2.1 Experimental Details and Data Acquisition SystemExperiments were conducted on a commercial LPBF machine (EOS M270) at NIST. The machine was integrated with three types of sensors, namely, a SWIR lens thermal camera (1350 – 1600 nm), high-speed video camera (4000 frames per second), and a photodetector. The SWIR thermal camera and photodetector capture the thermal aspects of the meltpool, whereas the high- speed video camera captures its shape and surrounding spatter pattern.Data was acquired at a sampling rate of 1 MHz. Figure 2 shows the schematic implementation of the setup.

Figure 2 shows the schematic implementation of the setup. The detailed explanation of the setup is available in [21, 22]. The test artifact, which is made from Nickel alloy IN625 (58% Ni, 23% Cr), has a steep overhang of . In this work, the sensor information was not captured continuously during the duration of the build, but only for three layers, at 6.06 mm, 7.90 mm, and 9.70 mm. The overhang is also sintered in these layers, rather than the solid portion alone. The process parameters are shown in Table 1. The overarching aim is to distinguish the sensor patterns that emerge during printing of overhang (2 hatches from the edge) and non-overhang features of the part shown in Figure 3.𝑜𝑜40.5Figure 2: Schematic layout of sensors installed on the LPBF 3D printer.1268The part is sintered per the stripe pattern shown in Figure 3(c and d); hence the laser sinters along the overhang four times (called stripes) for each layer past 4 mm build height.

The challenge is that the length of time in scanning the overhang portion progressively increases with the build height; therefore, the temperature trends recorded during printing the overhang will progressively increase with the build height.

In other words, the process exhibits a nonstationary trend, which limits the use of parametric statistical approaches, such as control charts, which rely on fixed distribution characteristics.Figure 3: (a) The part schematic (all dimensions in mm, drawings are not to scale) measuring 16 mm on all sides with 40° overhang, (b) as-built without supports, (c and d) Side-view and top views of the stripe pattern at the build height of 7.9 mm in the context of the thermal camera positionTable 1: Printing parameters used for fabrication of overhanging structurePrint Parameter Hatch distance Stripe width Stripe overlap Layer thickness Scan speedLaser powerValue 0.1 mm 4 mm 0.1 mm 20 μm 800 mm/s 195 W (infill) 100 W (pre-contour) 120W (post-contour)12692.2 Statistical Analysis of the Data(a)Image ProcessingThe aim of this section is to apply image processing and statistical signal analysis approaches to differentiate between OH and NOH printing conditions.

The intent is to isolate the sensor signatures including melt pool and spatter pattern during different build conditions - overhang (OH) and non-overhang (NOH) edge - from the thermal and high-speed video camera. A series of steps involving popular image processing procedures is implemented.Step 1: Each image frame was converted to a binary black and white image; the threshold value was selected heuristically and remains fixed. The image obtained after binarization is represented as𝐼𝐼𝑏𝑏Step 2: The aim is to isolate specific aspects of the meltpool, such as its size and spatter pattern. . As However, these are difficult to obtain without loss of information from the binarized image specific to the meltpool shape and the spatter pattern by a result, we create a binary mask following the procedure summarized herewith:𝐼𝐼𝑚𝑚Detecting the connected pixel components in the binary image; this is done using the matlab routine (imbinarize).

This routine determines whether, and to what degree a given pixel is connected to its nearest neighbors. The rationale is that a pixel in the meltpool region is well- connected to its neighbors (because the meltpool presents a larger contiguous area), compared to a pixel from the spatter pattern. Isolating the component (filtering away the components which have low connectivity). This leaves the largest area , i.e. , the meltpool shape, subsequently, and subtracting the meltpool results in the binary image with the spatter pattern. This image with the binary from . meltpool and spatter pattern is called the mask𝐼𝐼𝑏𝑏𝐼𝐼𝑏𝑏A Hadamard product (element-wise multiplication)yields images with the 𝐼𝐼𝑚𝑚 specific meltpool and spatter aspects retained. The resulting isolated meltpool and spatter for both the overhang and non-overhang conditions are shown in Figure 5.

The resulting isolated meltpool and spatter for both the overhang and non-overhang conditions are shown in Figure 5. pattern𝐼𝐼𝑚𝑚 ∙ 𝐼𝐼𝑜𝑜 = 𝐼𝐼𝑓𝑓𝐼𝐼𝑓𝑓1270𝐼𝐼𝑏𝑏Figure 4: Isolating the thermal signatures (melt pool and spatter pattern) using the sequence of image processing steps described.) for (a1 and a2) meltpool shape and (b1 and b2) spatter Figure 5: The isolated thermal frame ( pattern for the overhang and non-overhang build conditions, respectively. Note the elongated shape of the meltpool in (a2).𝐼𝐼𝑓𝑓Two statistical features are defined on the isolated pattern each column of plotting a histogram of pixel intensities in the isolated frame ( 𝐼𝐼𝑓𝑓 histogram count ( respectively.. The first is the sum of intensities for Σ). The second feature is based on ). This feature is called the. This feature is called column intensity ( 𝐼𝐼𝑓𝑓𝒥𝒥).

The first is the sum of intensities for Σ). The second feature is based on ). This feature is called the. This feature is called column intensity ( 𝐼𝐼𝑓𝑓𝒥𝒥). The procedure is shown schematically in Figure 6 (a) and (b) forand𝐼𝐼𝑓𝑓𝒥𝒥𝐽𝐽𝛴𝛴1271𝐽𝐽Figure 6: Two statistical features used for analysis of meltpool shape (a) column intensity (b) histogram countsShown in Figure 7 are the empirical cumulative distribution functions (ecdf) of the meltpool and for three overhang (OH) and non- spatter histogram counts ( overhang (NOH) build conditions.and column intensities (Σ𝒥𝒥)𝒥𝒥)Figure 7 (a1) and (a2) concern the meltpool histogram counts (and column intensities (respectively; the corresponding distributions for the spatter pattern are shown in Figure 7 (c) ) and (d). The following inference can be drawn from Figure 7. in Figure 7 (a1) cannot be visually distinguished.

The following inference can be drawn from Figure 7. in Figure 7 (a1) cannot be visually distinguished. for spatter is shown in Figure 7 (b1), and is able to discriminate𝒥𝒥)The ecdf of meltpool histogram counts ( The ecdf distribution of between OH and NOH conditions.The ecdf of meltpool histogram counts ( The ecdf distribution of between OH and NOH conditions.𝒥𝒥)Well-known statistical distributions, such as normal, exponential, Weibull, extreme value, Σ.𝒥𝒥𝒥𝒥Figure 7: Empirical cumulative distribution functions of thermal camera (a1 and a2) meltpool shape and (b1) and (b2) spatter pattern based on the histogram counts ( and column intensities (.Σ𝒥𝒥)𝒥𝒥)1272ΣFor instance, the probability plot for fitting the versatile Weibull statistical distribution on the spatter histogram counts ( is shown in Figure 8. In a good distribution the observations (histogram counts in our analysis) will fall within the confidence intervals.

In a good distribution the observations (histogram counts in our analysis) will fall within the confidence intervals. As observed from the probability plot, both tails of the ecdf deviates from the Weibull distribution; results for other distributions are worse. Hence, statistical analysis based on parametric distributions are found to be inadequate. This analyses were repeated with images obtained from the high-speed camera; the difference between the OH and NOH states was undistinguishable.𝒥𝒥)Figure 8: Probability plot for fitting Weibull distribution to thermal camera histogram counts for spatter pattern(b)Time-Frequency Analysis of the Photodetector Signal.The spectrogram is a widely used tool to capture the time-frequency and amplitude aspects of a signal simultaneously [23]. To obtain the spectrogram for the photodetector data, it is necessary to align the signal with the thermal camera frame (as the position information is not available).

To obtain the spectrogram for the photodetector data, it is necessary to align the signal with the thermal camera frame (as the position information is not available). A spectrogram is essentially a short time Fourier transform (STFT) taken over several windows of a signal. In other words, the thermal camera data is used as the so-called ground truth for distinguishing between the OH and NOH build conditions. As mentioned before, one frame of thermal camera translates to 555 data points (~ 555 microseconds) for the photodetector. As in the thermal image analysis in Sec. 2.2(a), the spectrogram analysis for two representative OH and NOH signals are shown in Figure 9.1273Figure 9: Spectrogram of the photodetector signal in (a) overhang and (b) non-overhang printing conditions for one layer.

Note that there is no discernable difference in the shape and magnitude of the frequency spectrum.Although a visual comparison between the overhang (OH) and non-overhang (NOH) spectrogram issues no significant difference, the patterns were statistically analyzed. Figure 10 shows the aggregate power spectral density (PSD) scatterplot in three different frequency bands of the photodetector signal, namely, 0-100 KHz, 200-300 KHz and 400-500 KHz for the overhang (OH) and non-overhang (NOH) build condition.A 2-sample t-test was performed to assess the hypothesis that the was no significant difference in mean PSD between the OH and NOH build conditions at each of the frequency bands. This hypothesis was rejected for the 400-500 KHz frequency band at the 95% confidence level; i.e., there is indeed a statistical difference in the PSD for the OH and NOH cases.

This hypothesis was rejected for the 400-500 KHz frequency band at the 95% confidence level; i.e., there is indeed a statistical difference in the PSD for the OH and NOH cases. Nonetheless, a statistical difference does not guarantee high classification fidelity, i.e., concluding the state of the build (OH vs. NOH) given a PSD is difficult due to the high variability as evidenced by the scatter plot in Figure 10.Figure 10: Scatter plot of power spectral densities at three different frequencies (0 – 100 KHz, 200-300 KHz and 400-500 KHz) with the p-value related to the 2 sample t-test.12742.3Spectral Graph Theoretic AnalysisThe aim of this section is to apply the spectral graph approach described in Sec. 2 to discriminate between the overhang and non-overhang build conditions The approach is compared against seven other popular machine learning approaches. For brevity, the parameter settings are encapsulated in the Appendix.

The data is for 250 randomly chosen sequences from each build condition.Predicted Build Conditions s e c o r Pl a u t c An o i t i d n o CThermal CameraVisible CameraPhotodetectorOH NOH OH NOH OH NOHOverhang (OH) 250 24 250 83 157 9Non-Overhang (NOH) 0 226 0 167 93 2413 Conclusions and Future WorkThis work proposed a spectral graph theoretic approach for monitoring the build status in laser powder bed fusion (LPBF) additive manufacturing (AM) process via a heterogeneous sensing array consisting of a photodetector, SWIR thermal camera, and high-speed video camera. The central idea of the approach is to convert the sensor data into a lower dimensional manifold, specifically, a weighted and undirected network graph. The intuition is that the topology of the graph is a surrogate indicator of the build states.

The intuition is that the topology of the graph is a surrogate indicator of the build states. Based on this premise, we tested the hypothesis that by quantifying and learning the signal patterns embedded in the topology of the network graph in terms its Laplacian Eigenvectors, process phenomena indicative of the build status is detected1275with higher fidelity compared to conventional machine learning approaches.

The specific conclusion is as follows.The proposed graph theoretic approach was applied to the sensor data with the intent to distinguish between the OH and NOH build states, the F-score obtained in the region of 80 to 95%, contingent on the type of sensors: F-score ~ 95% from the SWIR thermal camera  F-score ~ 83% for the high-speed video camera  F-score ~ 79% for the photodetector sensor.What other different types and more relevant microstructure-level defects, such as powder contamination, poor fusion, porosity delamination, etc., may be detected?What is the link between specific defects and sensor signal patters? In other words, is there a one-to-one link between a type of defect and its severity, and the sensor signature it manifests?What is the detection lag; does the detection accuracy improve with sensor redundancy?AcknowledgementsThe authors acknowledge the contribution of Dr. Brandon Lane and Dr.

Brandon Lane and Dr. Jarred Heigel of the National Institutes of Standards and Technology (NIST), United States Department of Commerce; the data for this work was generated at NIST. Dr. Lane and Dr. Heigel do not appear on the list of authors of the article in the conference proceedings pending review of its scientific merits and subsequent approval by NIST’s editorial review board per established policy.1276References[1] I. Gibson, D. W. Rosen, and B. Stucker, Additive manufacturing technologies: rapidprototyping to direct digital manufacturing. Springer, 2010.[2] J. z. Jacobsmuhlen, S. Kleszczynski, D. Schneider, and G. Witt, "High resolution imaging for inspection of Laser Beam Melting systems," in 2013 IEEE International Instrumentation and Measurement Technology Conference, 2013, pp. 707-712: IEEE.[3] S. Kleszczynski, J. Zur Jacobsmühlen, J. Sehrt, and G.

I. Shuman, B. Ricaud, and P. Vandergheynst, "A windowed graph Fourier transform," in2012 IEEE Statistical Signal Processing Workshop (SSP), 2012, pp. 133-136: Ieee.[16] O. DeJesús and M. T. Hagan, "Backpropagation through time for a general class of recurrent network," in 11th International Joint Conference on Neural Networks, Washington, D.C., 2001, pp. 2638-2642, Washington, D.C.[17] W. King, A. Anderson, R. Ferencz, N. Hodge, C. Kamath, and S. Khairallah, "Overview of modelling and simulation of metal powder-bed fusion process at Lawrence Livermore National Laboratory," Materials Science and Technology, 2014.[18] J. D. Lee and M. Maggioni, "Multiscale analysis of time series of graphs," in Proceedings of The 9th International Conference on Sampling Theory and Applications, Singapore, 2011.[19] X. Zhu and M.

D. Lee and M. Maggioni, "Multiscale analysis of time series of graphs," in Proceedings of The 9th International Conference on Sampling Theory and Applications, Singapore, 2011.[19] X. Zhu and M. Rabbat, "Graph spectral compressed sensing for sensor networks," in IEEE International Conference on Acoustics, Speech and Signal Processing (ICASSP), Kyoto, Japan, 2012, pp. 2865-2868: IEEE.[20] p. f. G. S. P. Workshop, in Graph Signal Processing Workshop, Philadelphia, PA, 2016.[21] B. Lane et al., "Thermographic measurements of the commercial laser powder bed fusionprocess at NIST," Rapid prototyping journal, vol. 22, no. 5, pp. 778-787, 2016.[22] B. Lane, E. Whitenton, and S. Moylan, "Multiple sensor detection of process phenomena in laser powder bed fusion," in SPIE Commercial+ Scientific Sensing and Imaging, 2016, pp. 986104-986104-9: International Society for Optics and Photonics.[23] S. Qian and D. Chen, Joint time-frequency analysis: methods and applications.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIn-Process Monitoring of Cross Contamination in Laser Powder Bed Fusion (L-PBF) Additive Manufacturing (AM)M. Jamshidinia, P. Boulware, J. Marchal, H. Mendoza, L. Cronley, S. Kelly, S. Newhouse EWI 1250 Arthur E Adams Dr., Columbus, OH 43221AbstractCross contamination in laser powder bed fusion (L-PBF) Additive Manufacturing (AM) could locally change the chemical composition and stress distribution in a component. It also could result in the formation of flaws, and consequently lower the mechanical performance of a component. In this study, the in-process monitoring and detection of cross contamination was investigated in L-PBF process, also known as direct metal laser sintering (DMLS).

In this study, the in-process monitoring and detection of cross contamination was investigated in L-PBF process, also known as direct metal laser sintering (DMLS). A setup was designed and fabricated at EWI, where contaminant materials could be introduced on the powder bed without interrupting the fabrication process or breaking the chamber environment. Tungsten particles were used as the contaminant material, in the matrix of Inconel 625. Six levels of contamination were calibrated, and were introduced in two static and dynamic modes. Photodetector, spectrometer, and optical camera were used for the data acquisition. One of the sensors showed the most promising results.

Photodetector, spectrometer, and optical camera were used for the data acquisition. One of the sensors showed the most promising results. X-ray computed tomography (CT) and optical microscopy were used to validate data collected by the sensors.Keywords: cross contamination, in-process monitoring, spectrometer, photodetector,laser powder bed fusion (L-PBF)IntroductionThe layer-by-layer fabrication of metallic components in laser powder bed fusion (L- PBF) processes occurs as a result of interactions between a high power laser beam and metallic powder particles. Several studies have been done on the in-process monitoring of PBF processes. G. Zenzinger et al [2015] used a high-resolution camera system to obtain thermal images from the L-PBF process. A three-dimensional (3D) digital map was prepared from the structural irregularities and structural disorder detected by the optical tomography (OT) method.

A three-dimensional (3D) digital map was prepared from the structural irregularities and structural disorder detected by the optical tomography (OT) method. The 3D map was correlated to mechanical performance and failure analysis of specimens. T. Furumoto et al. [2012] used a high speed video camera to investigate the melting and solidification of a ferrous alloy during the L-PBF process. The shape of molten pool as well as the powder scattering influenced by process parameters was visualized. J. P. Kruth et al [2007] used a photodetector and a complementary metal–oxide–semiconductor (CMOS) to monitor and control the size of molten pool. Some coupons with overhanging features were produced with and without the feedback control system. Results showed that both sensors can be integrated in a feedback control system and the melt pool area can be controlled. S. Price et al.

Results showed that both sensors can be integrated in a feedback control system and the melt pool area can be controlled. S. Price et al. [2013] used a near infrared thermal imager for temperature measurements in the electron beam powder bed fusion (EB-PBF) of Ti-6Al-4V. The influence of overhanging features on the size of the molten pool was analyzed at various build heights. B. Foster et al. [2015] used a high-resolution DSLR camera to collect images under oblique illuminations of fused and pre-placed powder layers. Various types of build defects were detected. Also, a three dimensional (3D) reconstruction of a1366fabricated part and potential defects was produced using the high exposure optical images. S. Kleszczynsk et al. [2012] used a high-resolution imaging camera to detect errors caused by predefined process irregularities during the L-PBF of Inconel 625. A monochrome charge coupled device (CCD) was sued for this purpose.

A monochrome charge coupled device (CCD) was sued for this purpose. The same camera was also used to collect data from a cylinder produced with contaminated powders. Results showed a distinguishable difference between the clean and contaminated cylinders. However, the CCD camera did not provide any additional data about the level of contamination.So far, a limited study has been conducted on the monitoring and detection of cross contamination during the PBF process. In this study, the monitoring and detection of cross contamination were investigated, where the main focus was to provide an experimental and qualitative comparison between the collected data from three sensors, and develop a relationship between the data and the contamination levels.

The quantitative analysis of data such as correlating a signal characteristics to contamination levels and contaminant materials will be conducted in the future study.Design of Experiment MaterialsIn this study, a rectangular specimen with a 10×10×15.20-mm dimension was fabricated with Inconel 625. Tungsten powder was the contaminant material to be deposited in between the Inconel 625 layers. A carbon steel plate was sued as the build platform. All powders used in the contamination study were virgin powders.The Cross Contamination SetupA setup was designed and fabricated by EWI, for the deposition of precise amount of contamination at each desired step. To minimize the possibility of undesired powder deposition during the fabrication process, the setup was tested for powder leakage. The entire platform under the contamination setup, including dispenser, build plate, and collector was covered. The test was conducted by executing the powder deposition for 100 times.

The entire platform under the contamination setup, including dispenser, build plate, and collector was covered. The test was conducted by executing the powder deposition for 100 times. Not a detectable amount of powder was leaked from the powder container on the build platform and collector.Deposition CalibrationFigure 1 shows the six levels of contamination used for tungsten respectively. There was a 0.5-mm gap between the contamination setup nozzle (on the bottom surface) and the powder bed top surface.The contamination levels were categorized in two groups. In the first group containing L1, L2, and L3, the contaminant powder was deposited while the recoater was stopped (static category). As a result, a spot-shaped profile of contaminant powder was formed. In the second category (L4, L5, and L6), the contaminant powders were deposited, while the recoater was moving (dynamic category). So, the contaminant powders were spread as an1367Figure 1.

In the second category (L4, L5, and L6), the contaminant powders were deposited, while the recoater was moving (dynamic category). So, the contaminant powders were spread as an1367Figure 1. The six levels of contamination deposition for tungsten. Lc and Wc refer to the length and width of a contamination profile.Figure 2. The unit volumes of deposited powders for the six contamination levels for the tungsten (W) powder. L1 to L3 correspond to the static powder deposition. L4 to L6 correspond to the dynamic powder deposition.elongated line and formed a linear profile on the powder bed. Figure 2 plotted the deposited unit volumes of powder corresponding to the contamination levels.Build LayoutTable 1 shows the build layout used for the contamination study. The contaminant powder was deposited on a single layer, while 19 layers of baseline parameters were deposited below and above it.

The contaminant powder was deposited on a single layer, while 19 layers of baseline parameters were deposited below and above it. Baseline parameters refer to layers of Inconel 625 that were produced without the introduction of contaminant powders. In order to investigate repeatability of data acquisition, the whole set of six contamination levels was repeated three times, consecutively. Data acquisition was conducted for a total of 10 layers per each contamination layer. Two layer before the contamination, the contamination layer, and seven layers after the contamination were recorded. It should be noted that the labeling of layers on data acquisition data was different than the actual layers numbering.1368Table 1. the build layout for the contamination study.

It should be noted that the labeling of layers on data acquisition data was different than the actual layers numbering.1368Table 1. the build layout for the contamination study. The contamination set of L1 to L6 was deposited three times, and was followed by the cover layers.Contamination Set #Base Line (BL) / Contamination Layer (Ln, n= 1 to 6)Start LayerEnd LayerStart Height (mm)End Height (mm)Contamination Layers (×3 iterations)BL L1 BL L2 BL L3 BL L4 BL L5 BL L6 BL1 20 21 40 41 60 61 80 81 100 101 120 36119 20 39 40 59 60 79 80 99 100 119 120 3800.04 0.80 0.84 1.60 1.64 2.40 2.44 3.20 3.24 4.00 4.04 4.80 14.440.76 0.80 1.56 1.60 2.36 2.40 3.16 3.20 3.96 4.00 4.76 4.80 15.20Cover LayersFigure 3. A rectangular specimen made of Inconel 625, containing contamination layers.Figure 3 shows the rectangular specimen made of Inconel 625, with embedded layers of contamination.

A rectangular specimen made of Inconel 625, containing contamination layers.Figure 3 shows the rectangular specimen made of Inconel 625, with embedded layers of contamination. Optimum process parameters were developed for the fabrication of Inconel 625 in EWI’s test bed.Data AcquisitionPhotodetector, spectrometer, and optical camera were used for the data acquisition. An Ocean Optics HR2000+ sensor was used as the spectrometer. The detection range was 200 to 1100 nm, with an exposure set at 1 ms (frequency of 1000 Hz). The photodetector was a PDA36A sensor from Thor Labs, with a detection range was 350 to 1100 nm. The gain was set at136940dB. Also the signal amplification was tuned to maximize the signal without saturation. A Flea3 optical camera from Point Grey was used to record snapshots of the powder bed top surface after the deposition of the contaminant powders as well as after the fusion by laser.

A Flea3 optical camera from Point Grey was used to record snapshots of the powder bed top surface after the deposition of the contaminant powders as well as after the fusion by laser. The camera had an 8.8 mega pixel resolution, with a field of view of 1.0 × 0.58 in, and was able to capture one frame per second. Both of the spectrometer and photodetector were installed on the optical table and aligned with the on-axis signal.Figure 4 shows the results of data acquired from a tungsten-contamination build. The two graphs on the top and the middle plotted data collected by the spectrometer and photodetector, respectively. The two images shown in the bottom row of each contamination level were taken after the deposition of the contaminant powder and after the fusion by laser. The darker profiles on the optical images refer to the tungsten powder spread over the powder bed.Figure 4.

The darker profiles on the optical images refer to the tungsten powder spread over the powder bed.Figure 4. Results of data acquisition using spectrometer (top), photodetector (middle), and optical camera (bottom) during the deposition of tungsten contaminants. L1 through L6 refers to the six levels of contamination. Also, 2 refers to the second iteration of contamination deposition.1370SpectrometerRegardless of the contamination level, the data collected by the spectrometer showed signals (peaks) at identical wavelengths. However, the signals had different intensities (strengths). A general relationship could be established between the contamination levels and the maximum signal strengths. Spectrometer detected a stronger signal with an increase in the compactness (unit volume) of a contamination level (Figure 4). For example, L6 showed a signal with the maximum strength.

Spectrometer detected a stronger signal with an increase in the compactness (unit volume) of a contamination level (Figure 4). For example, L6 showed a signal with the maximum strength. However, the contamination levels with low unit volumes, such as L1, L4, and L5 did not show signals with distinguishable differences. It could be related to the limited capability of the spectrometer in detecting contaminations with low unit volumes. Therefore, only the tungsten contamination levels with a unit volume of 0.006 mm3/mm2 or higher could be distinguishably detected by the spectrometer used in this study.It should be noted that the spectrometer used in this study had a data acquisition frequency of 1000 Hz.

A spectrometer with a higher frequency might be able to capture more accurate data from different contamination levels that could make collected data more distinguishable.Optical cameraAccording to Figure 4, all contamination levels except L1 and L4 showed clearly detectable discolorations on the top surface of the specimen after the deposition of contaminant powder. In some cases such as L3 and L6, the discoloration was detectable even after the fusion of the powder bed. Because the unit volumes deposited in L1 and L4 were low, not enough contrast were formed between the contaminated zone and rest of the top surface.

Because the unit volumes deposited in L1 and L4 were low, not enough contrast were formed between the contaminated zone and rest of the top surface. The optical camera used in this study was able to detect the tungsten contamination levels with a unit volume of 0.005mm3/mm2 and higher.It should be noted that the development of an improved illumination system, as well as the synchronization of the optical camera with an image analysis code would improve the quality and reliability of data captured by the optical camera.PhotodetectorThere was a strong relationship between the contamination levels and the data collected by the photodetector. Figure 5 illustrates the characteristics of signals generated by the photodetector. The horizontal axis shows the data acquisition time. Each set of data is the representative of one hatch line (Figure 5-a).

The horizontal axis shows the data acquisition time. Each set of data is the representative of one hatch line (Figure 5-a). Depending on unit volume deposited on the powder bed as well as the profile of the deposited contaminant powder, a single peak (Figure 5- b) or multiple peaks (Figure 5-c) were recorded. Figure 5-d shows the profile of the deposited contaminant powder for the contamination level of L3. Lc and Wc refer to the length and width of the deposition profile, respectively. The dashed-black arrow shows the recoating direction.Figure 6 illustrates the contamination profiles with their associated photodetector data. Any of the images consisted of data collected from six consecutive hatches and was equivalent to1371Figure 5. The characteristics of signals generated by the photodetector, a) each data set was a representative of one hatch line.

The characteristics of signals generated by the photodetector, a) each data set was a representative of one hatch line. B) Single peak, c) Multiple peaks, d) The distribution of tungsten powder on the powder bed, where Lc and Wc refer to the length and width of the contaminant powder profile, respectively. The black dashed line shows the recoating direction.Figure 6. Correlation between the contamination profile and the photodetector data. The number and strength of signals detected by the photodetector varied with contamination levels. The black arrow shows the laser scanning direction as well as the recoating direction.0.1 seconds of the scanning process. The data collected from the six contamination levels were different in the following items:Number of peaks per hatch  Intensity of the peaks  Number of hatches with high intensity peaks1372The black arrow in Figure 6 shows the laser scanning direction as well as the recoating direction.

At a contamination level with a longer length (Lc), the contaminant powder was exposed to the laser beam for a longer period of time. As a result, multiple peaks were generated and recorded by the photodetector. Therefore, the presence of multiple peaks per hatch at L3-2 and L6-2 could be related to the elongated length of the deposition profile.The intensity of a peak could be related to deviation of the molten pool from a default condition. The application of a contaminant powder profile with a higher compactness (or a higher unit volume) results in a larger deviation from a default condition. Therefore, even though both of the L5-2 and L6-2 contamination levels had an elongated deposition profile and showed multiple peaks per hatch, L6-2 had stronger peaks were due to its higher compactness (unit volume) of contaminant powder.More hatches with strong peaks could be recorded if a contamination profile had a wide contamination width (Wc).

However, a high compactness was also necessary for the formation of multiple hatches with high intensity peaks.Detectability after the contaminationThis section investigated the possibility of detecting tungsten contaminations in consecutive layers. Because the photodetector showed the most promising data so far, data collected by the photodetector were analyzed for this purpose. Figure 7 compares the photodetector data collected from four consecutive layers of each contamination levels. The first set of data on the top corresponds to the contamination layers, where a contaminant powder was deposited on it. This layer was identified by n. Layer n+3, at the bottom row, represents the third layer on top of a contamination layer. As shown, almost all contamination levels still showed some distinguishable signals on the second layer (n+1); however the signals were weaker than those on the first layer.

As shown, almost all contamination levels still showed some distinguishable signals on the second layer (n+1); however the signals were weaker than those on the first layer. However, by the deposition of two more layers, most of the initially detected signals disappeared. Even at L3, with a high unit volume deposited on the powder bed, almost all signals were faded out on layer n+3. L5 and L6 were the only two contamination levels that still showed one signal on layer n+3. These results suggest that a contamination with tungsten, even at very high levels, could only be detected for up to three layers (120 µm) on top of where the contaminant powders were introduced. It has to be noted that none of the contamination levels showed any signals on layer n+4.Post process inspection Metallurgical analysisThe specimen with embedded tungsten contaminant was sectioned, mounted, and polished, for further metallurgical analysis.

An alcohol-based Kalling’s etching solution was used for the primary etching. Then specimens were etched using a 10% chromic acid solution at 2.4 volts. Figure 8 shows the low magnification (25X) images of the cross sections. The white arrows indicate the contamination deposition layers. The black arrow indicates the buildup direction. The tungsten particles with a white color could be clearly detected on the matrix of Inconel 625.1373Figure 7. Data collected from photodetector from the six levels of contamination with tungsten. Four layers per contamination level are studied.1374Figure 8. Low magnification optical images of the specimens contaminated by tungsten. The white arrows indicate the contamination deposition layers. The black arrow indicates the buildup direction.Figure 9. High magnification optical images of the specimens contaminated by tungsten. The white arrows indicate molten pools with distorted appearance.

High magnification optical images of the specimens contaminated by tungsten. The white arrows indicate molten pools with distorted appearance. The black arrow indicates the buildup direction.Figure 9 shows a high magnification image of the contaminated specimen. The solid black arrow indicates the buildup direction. Similar to Figure 8, the white tungsten (W) particles could be observed with a high contrast with the Inconel 625 matrix. These cross sectional images could also provide some additional information about the L-PBF process. The dashed-line (1) indicates the shape of a molten pool, penetrated into the previously deposited layer. Also, the overlaps between tracks could be recognized. However, the layer below the indicated molten pool had a different appearance, with no indications of molten pool profiles. The dashed line and circle (2) shown at the bottom of the tungsten specimen are representative of the hatching directions in the two consecutive layers.

The dashed line and circle (2) shown at the bottom of the tungsten specimen are representative of the hatching directions in the two consecutive layers. When the laser scanning direction is normal to the polishing surface, molten pools are formed normal to the polishing surface and could be observed on etched specimens. Because the hatching direction was rotated 90 degree between1375layers, the molten pool of the consecutive layer formed parallel to the polishing surface, and was observed as a long continuous stripe.X-Ray Computed Tomography (CT)Metallurgical analysis provided some information about the contaminated specimens. However, it was limited to only one cross section per specimen. To better understand the distribution and melting of contaminant powders, one specimen per contaminant material was inspected using X-Ray Computed Tomography (CT). Figure 10 illustrates the 3D reconstruction of the specimen contaminated with the tungsten powder particles.

Figure 10 illustrates the 3D reconstruction of the specimen contaminated with the tungsten powder particles. The three planes represent the three axes of the Cartesian coordinate system. The recoating direction is shown by the black arrow. The vertical cross-sectional views of the reconstructed specimen are presented in Figure 11. The white arrow indicates the recoating (or powder spreading) direction. Seventeen layers of contaminated powders could be detected, while the second layer from the bottom (Level L2) was missing. The missing layer was indicated using the white stars in Figure 11. Figure 11-a shows the cross sectional view normal to the recoating direction. The white-dashed line box indicates the second set of contamination levels, from L1 to L6, embedded in the specimen. The results of the CT confirmed that the minimum and maximum amounts of powder were deposited in L4 and L6, respectively.

The results of the CT confirmed that the minimum and maximum amounts of powder were deposited in L4 and L6, respectively. Figure 11-b provided additional information on the tungsten powder distribution, with a cross-section parallel to the recoating direction. The three contamination levels of L6 were shown to distinguish the three consecutive sets of contamination levels. The contamination levels L4 through L6 had elongated powder tails, due to their dynamic powder deposition system. L1 and L2 with the static powder deposition had the shortest tails of powder. The contamination level L3 was expected to have a short tail, as well. But a longFigure 10. Three-dimensional (3D) reconstruction of the specimen contaminated with tungsten powder particles. The black arrow shows the powder recoating direction.1376Figure 11.

But a longFigure 10. Three-dimensional (3D) reconstruction of the specimen contaminated with tungsten powder particles. The black arrow shows the powder recoating direction.1376Figure 11. The vertical cross sectional views of the specimen contaminated by the tungsten powder particles, a) cross section normal to the recoating direction, b) cross section along the recoating directionpowder tail was observed on all of the three iterations of L3 levels. The unexpected powder distribution in L3 was attributed to the transfer of contamination between consecutive layers, which will be studied in future works.Figure 12 illustrated the horizontal view of the tungsten powder distribution in the six contamination levels. There was a noticeable match between the calibration images shown in Figure 1 and the actual distribution inside the specimen.1377Figure 12. The horizontal cross sectional views of the specimen contaminated with the tungsten powder.

The horizontal cross sectional views of the specimen contaminated with the tungsten powder. Images show the six contamination levels, along the buildup direction (shown by the white arrow).ConclusionThe cross contamination of Inconel 625 was studied during the L-PBF process. A setup was designed and fabricated by EWI for this purpose. The Tungsten powder was used as the contaminant material. Three sensors including spectrometer, photodetector, and optical camera were used to monitor the process. The contaminated specimens were studied metallurgically. Also, X-ray CT was used to reconstruct the 3D model of the contaminated specimens and investigate the distribution of contaminant powders.The results of this study could be summarized as following:1. Among the current sensors used in this study, photodetector was the most promising sensor that detected all of the contamination levels of tungsten.

Among the current sensors used in this study, photodetector was the most promising sensor that detected all of the contamination levels of tungsten. Spectrometer and optical camera were able to detect the contamination levels with a unit volume equal to or higher than 0.006 mm3/mm2 and 0.005mm3/mm2, respectively.2. Even at the highest level of tungsten deposition, a contamination could be detected for up to three layers (120 µm) on top of where the contaminant powders were initially introduced.13783. The size of area contaminated area became larger in consecutive layers, due to the transfer of unfused particles along the recoating direction, as well as the buildup direction.4. The minimal dilution of the tungsten particles in the matrix could be explained by considering the melting temperatures of materials used in this study. The process parameters were optimized for the L-PBF of Inconel 625, which was too low for the melting of tungsten.

The process parameters were optimized for the L-PBF of Inconel 625, which was too low for the melting of tungsten. Future StudyThe detection of contamination in L-PBF processes could be further developed throughthe implementation of the following items:1. Detectability of other materials will be investigated. 2. The quantitative analysis of data including identifying types of contaminating materials will be conducted in the future study. The application of a spectrometer with a higher data acquisition frequency could help in analyzing spectra and elements.3. Application of the following sensors could provide additional information from the molten pool:a. High speed charge coupled device (CCD) camera b. Local thermal sensor c. Global thermal sensor (IR camera)4.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIN-PROCESS UV-CURING OF PASTY CERAMIC COMPOSITEStefan Mischliwski1, Matthias Weigold11 Institute of Production Management, Technology and Machine Tools (PTW),Technische Universität Darmstadt, Darmstadt, GermanyAbstractWithin recent years,a wide range of additive manufacturing processes have been developed. While powder bed based fusion processes like selective laser melting and melting processes like fused layer modelling are being increasingly used in industrial applications, prototyping other processes are in the initial stage. This paper develops a new method for an extrusion-based process of pasty UV-curing ceramic composite material. The method proposes an approach to continuously cure the material while it is deployed to reduce process time and generate complete cured parts.

The method proposes an approach to continuously cure the material while it is deployed to reduce process time and generate complete cured parts. A milling machine has been modified with a syringe and a UV-light source to accommodate the process. Experimental studies have been carried out to examine the influence of the process parameters on the curing process. As a result, a parameter set has been found to make fully dense and cured ceramic composite parts.IntroductionAdditive manufacturing (AM) processes are currently the base for a variety of research projects. This emerging field in manufacturing technologies functions on the common principle of building up solid parts directly from 3D-CAD data by adding material layer by layer. Metal, plastic, ceramic or, as expected in the future, human tissue, can be used in AM. Particularly in areas like medical and dental technology, extensive requirements are placed on the production processes and resulting products.

Particularly in areas like medical and dental technology, extensive requirements are placed on the production processes and resulting products. Using new production processes, the main goal is to match the specific material properties in the conventional production process. This is possible through the controlled use of a variety of process parameters. Stereolithography is the state-of-the-art in the field of AM of UV- curing ceramic composites. However, this process can only be used on flat surfaces so that it is currently not possible to print on freeform surfaces. Therefore, a new process must be developed to apply pasty material onto freeform surfaces. Precautions must be considered in order to stabilise the location of the material application and to prevent the material from running.This paper investigates the in-process UV-curing concept for an extrusion-based 3D printing process of pasty ceramic composite.

To do so, a state-of-the-art conventional UV-curing process within a UV oven is investigated and critical factors are identified. Adjusting the hardware setup, the process is then ready to be transferred into a different environment in the form of an extrusion-based 3D printer. The influence of critical factors is examined and process parameters are adjusted with a focus on reaching the same surface hardness as the curing process within a UV oven.The UV-curing process of composites themselves has been studied for years so that many interactions of individual process parameters are known. However, these have yet not been investigated for curing during an extrusion-based material application. For this purpose, the interactions must be re-examined and assessed from the perspective of the overall process.1707State-of-the-artMany correlations of process parameters of the curing process have already been outlined in studies.

It has been shown that there is a partially linear relationship between the surface hardness of the cured ceramic composite and the light intensity used. However, this linearity only applies to a limited area [1]. A reduction of the light intensity used can be compensated by an increase in the exposure time [2]. Another insight is the decreasing hardness depending on the material depth or sample height. This can be explained by the absorption of light entering the translucent material. For this purpose, a model was created, depending on the depth of cure, light intensity and curing time. However, due to a maximum achievable surface hardness, the model can only be used in a limited range [3]. In a further study to determine parameters relevant to the curing depth, in addition to the curing time and the light source, the material composition was also identified [4].

In a further study to determine parameters relevant to the curing depth, in addition to the curing time and the light source, the material composition was also identified [4]. It turned out that in all examined composites, the depth of cure remains almost constant to a certain depth, between about 0.5-5 mm at a lighting duration of 300 s and then decreases rapidly. It has also been shown that as the illumination duration decreases, the cure depth decreases as well, and that the proportion of starters in the different composites is crucial for the curing depth. In [5], the material composition, the light intensity of the light source, the curing time and the distance between the tip of the light source and the body to be cured were identified as the four crucial factors of the curing process.

In addition to the affirmation that lighter and more translucent materials allow a greater hardening depth under the same conditions, the result was a correlation with increasing distance from light source to the body to be cured. As the distance increases, a significant reduction in polymerisation occurs at >20 mm. To determine the influence of the light source, [6] compares conventional halogen lamps with an LED light source. Both achieve a more than sufficient hardening depth.Research methodologyVITA VM LC flow (Enamel), a commercially available ceramic composite for the top layer of dental applications, is used as raw material to conduct the experiments. In dental laboratories, a dental technician manually applies the material according to the processing instructions given by the material manufacturer VITA Zahnfabrik H. Rauter GmbH (VITA). The part is than placed in a VITA recommended UV oven for the UV-curing process to reach its necessary surface hardness of > 60 HV0.5.

Rauter GmbH (VITA). The part is than placed in a VITA recommended UV oven for the UV-curing process to reach its necessary surface hardness of > 60 HV0.5. In this study, the bre.lux Power Unit 2 UV oven is used. This curing process serves as a reference during the experiments. The light source in the UV oven contains 72 LEDs, the exact type and wavelength spectrum of which is unknown. The ceramic composite Enamel consists of an organic matrix, a disperse phase and a composite phase. Monomers, in this case, di-acrylates, form the organic matrix which react with one another by radical polymerisation and cure. To start the curing process, the organic matrix is enriched with two initiators:(cid:120) Kampherchinon with an initiating wavelength of 468 nm (cid:120) TPO with an initiating wavelength of 380 nm.Colour pigments are another ingredient, but just for cosmetic reasons. The disperse phase consists of silica and zirconium dioxide.

The disperse phase consists of silica and zirconium dioxide. These are responsible for hardness and polishability after curing. The composite phase represents the functional connection between organic matrix and disperse phase.1708As none of the above studies deal with the ceramic composite Enamel used here or with the goal to integrate the curing process into an AM process, the individual factors must be considered again.The curing process with the oven used fulfills the requirements of DIN EN ISO 10477:2017 [7]. To better understand the process, hardware components of the oven are examined. Here, the wavelength spectrum of the LEDs used and the light intensity over time are of particular interest. The LEDs of the oven use the wavelengths 380 nm and 468 nm to initiate the hardening process of Enamel. Because the LED wavelength bandwidth is narrow in comparison to other light sources, it is assumed that intensity peaks are close to these initiation wavelength [8].

Because the LED wavelength bandwidth is narrow in comparison to other light sources, it is assumed that intensity peaks are close to these initiation wavelength [8]. To carry out the wavelength measurement, a hand spectrometer, a pulse generator, a specially designed mounting box for the LEDs and optical Mirror and filter are used and arranged according to Figure 1. The LEDs of the UV oven are mounted in groups of three on PBC printed circuit boards. One printed circuit board is placed in a specially made box, which allows the shielding of the individual LED light beams, whereby the control technology of the UV oven for driving the LEDs could continue to be used. In order to take the temperature dependency of LEDs into account, the measurement was carried out over 720 sec, the reference time specified by VITA.

In order to take the temperature dependency of LEDs into account, the measurement was carried out over 720 sec, the reference time specified by VITA. Therefore, the control technology of the UV oven runs an internal programme of 360 sec twice.uv- / bluelightmirroruv-ovenremoved LEDs[:Ji----hand spectrometerpulse-generatorFigure 1: setup for the measurement of LED wavelength spectrumIn a further step, the light intensity is measured over the reference time of 720 sec. The maximum light intensity is selected to achieve maximum curing and curing depth during the limited curing time. For this measurement, it is important to keep the functionality of the oven’s LED control intact. By constantly introducing energy into the body, the temperature also has influence on the curing process.

By constantly introducing energy into the body, the temperature also has influence on the curing process. In the Enamel data sheet [9] has described that a temperature range between 60 °C and 80 °C is beneficial for the polymerisation.In order to implement a curing process into a 3D printer, a light source has to be available inside the interior of the printer. For this purpose, a holder for numerous LEDs is developed in which different LED types can be installed. In order to create a possibility for transporting the light from the source to the body, the PURAVIS® GOF70 light guide from SCHOTT AG is used for the required wavelengths due to the transmission rate. With this hardware setup, preliminary tests are carried out. In these tests, the surface hardness of the top, the bottom and the ratio of these two are measured and compared. These represent the quality feature according to which the quality of the process can be assessed.

These represent the quality feature according to which the quality of the process can be assessed. For comparison, the experiment is also carried out in the UV oven.1709The test body is manufactured manually according to DIN EN ISO 10477:2017 [7] specifications. The test body is a 1 mm thick disc with a diameter of 15 mm. The hardness measurement on the surfaces is carried out after Vickers.An LED pairing of suitable wavelengths, the influence of the light intensity, the curing temperature and the number of light sources are examined in respective test series. The post-curing time, the time after the LEDs turn off until the surface hardness is measured, is standardised to 20 min because of results reached in [10]. Here, no significant differences in surface hardness are shown in the time windows between 20 and 60 minutes or one and seven days post-curing.

Here, no significant differences in surface hardness are shown in the time windows between 20 and 60 minutes or one and seven days post-curing. The results obtained are used to draw conclusions on how a further optimisation can be designed to reduce lighting time and therefore the curing process time. The main objective is to achieve the same or better component properties, so it is sufficient to look at the process qualitatively. Since the curing is taking place as a radical polymerization, it needs to be considered that this is a chemical reaction over time. All experiment are carried out in an air-conditioned measuring room.Process developmentThe process development is carried out to estimate the potential of the UV-curing process for integration into an extrusion based additive manufacturing process.

The wavelength peaks of the three LED types are at 450 nm, 459 nm and 395 nm, as shown in Figure 2 a-c. Figure 2 d shows the wavelength spectrum of all three LEDs put together. Here, the wavelength to trigger the initiators for the curing process, Kampherchinon (468 nm) and TPO (380 nm), are also marked in the graph. It is shown that the intensity peaks of the used LEDs do not align with the initiator wavelengths. This indicates the potential for optimisation through the use of more suitable LEDs in which the intensity peaks are at the initiator wavelengths. Based on these results, a larger wavelength range is covered in the later LED selection with the attempt to ensure optimised performance.A second result of the hardware setup is the radiation intensity of the LEDs. It can be concluded from the results that the lighting in the oven runs a specific internal programme.

It can be concluded from the results that the lighting in the oven runs a specific internal programme. The radiation intensity starts with a jump to 50% of the possible power, is held there for about 30 s, and then increases linearly to 100% intensity. A slightly decreasing radiation intensity after approx. 80 sec is due to thermal effects within the LEDs themselves, as they show decreasing efficiencies with increasing temperatures. From 180 sec, the modulation programme of the UV oven begins with a variation of the radiation intensity between 50% and 100% for intervals of a few seconds.1710:;; ~ i!' ·;;; C:~r-- 0 .,. 0 00 "' "' "'Peak at 395n mN "' "' .,. r..: .,. .,. r-- "' ..... "' "' "'..... .,; "' .... ai u:i "' ..... .,. .,."1 C> "' r..: .,. ..... "' "'<rJ.,.<rJ000:;; ~ i!' ·;;; C: ~ .!:Peak at 450nm.,. .,.' ..... "' "' 0 .,; "' .... ai u:i "' .... .,. "' r-- "' ..... "' .,. r..: 0 .,. .,. "' "' "' "'r-- 00"' r..: .,.

. - - Initiat or wavelengt hat 468nm\ \I~466.6483.9501.05 17.9dwavelength [nm)Figure 2: a-c) wavelength spectrum of the three different LEDs used in the uv-oven; d) combined wavelength spectrum of all three LEDs used in the uv-oven and marking of the qualitative intensity of the initiator wavelengthMeasurement of surface hardness over lighting time For the development of a new curing process, a test setup is designed. In this setup, the variable combination of same or different LEDs is possible. A base body is designed to hold up to ten LEDs while ensuring sufficient heat dissipation for the LEDs. An upper body serves as a carrier of the light guide and is fitted appropriately. To carry out the experiment, individual LEDs are fastened in the base body and connected to a power source and a controller.

To carry out the experiment, individual LEDs are fastened in the base body and connected to a power source and a controller. The ends of the light guides are positioned 20mm above the samples to be cured, as shown in Figure 3.sampleDevice after: DIN EN ISO 10477:2017Figure 3: test sample with the light emitting end of the light guide1711During the experiments, LEDs with wavelength peaks shown in Table 1 will be used, wheretwice-occurring wavelengths differ by different power and technical specifications:Table 1: List of LEDs and wavelength usedUV LEDblue LEDNamewavelengthNamewavelengthLED A LED B LED C LED D370 nm 385 nm 395 nm 395 nmLED E LED F LED G LED H LED I445 nm 458 nm 458 nm 470 nm 475 nmIn the first series of experiments, the lighting time tB is varied between one and twelve minutes. The LED pair consisting of the UV LED B and the blue LED I was operated with a current of 0.5 A.

The LED pair consisting of the UV LED B and the blue LED I was operated with a current of 0.5 A. The graph in Figure 4 shows that, from an exposure time of more than 5 minutes, no significant influence on the total hardness HG was observed, taking into account the standard deviation. In order to exclude the lighting time as an influencing factor in the following series of experiments, lighting time tB is set to six minutes. The experiment is repeated with other LED pairings and show similar results.He [HVl] 34,032,031330,030,528,026,024,022,020,0123 ts [min] 45612Figure 4: Total hardness HG over the lighting time tB, LED pairing B & I, current IB = 0.5 A, post-curing time th = 20 min, curing temperature TA = 21 ° C, number of LEDs n = 2Measurement of surface hardness over different wavelength To achieve a process in which the two initiators are triggered optimally, the LEDs are paired with each other, consisting of a UV LED and a blue LED.

The goal is to find a combination that is particularly suitable for the process. For the experiments a lighting time of six minutes, a current of 0.5 A, and a post-curing time of 20 minutes are used. The 20 possible LED pairings are examined1712and the surface hardness of the prepared samples is measured at six points, three at the top and three at the bottom. The results are shown in Figure 5. Regardless of the influence of the five blue LEDs, the two UV LEDs A and B appear more suitable for the curing process than the LEDs C and D, since the achieved average surface hardness is higher. A statement about a specific blue LED could not be made with this measurement series because whether the results have a big range, e.g. LED E from 28 to 33.3 HV1, or the values are in too low a range, e.g., LED I from 29,6 to 31,2 HV1, additional iterations are necessary.HG [HVl) 34, 033,533,032, 031,030,029, 028,027, 027,327,026,0cii w (!J ciS ciS cc u u uu.

ciScii w u. ciS ciS C Cw ciS <t <t <t <t <tu. ciScii w ciS ccu. ciS cc~ J: ciS cc cc Figure 5: Total hardness HG of twenty LED pairs, lighting time tB = 6 min, current IB = 0.5 A,~ J: ciSJ: ciS u ucii J: (!J ciS ciS C C Cpost-curing time th = 20 min, curing temperature TA = 21 ° C, number of LEDs n = 2Measurement of surface hardness over operating current The radiation intensity of the light sources is examined as an influencing factor to the curing process. By varying the operating current, the radiation intensity is adjusted. These are in an approximately linear correlation to each other [11]. The temperature of the LEDs is kept constant during the experiments by active cooling. As a result, a decrease in the efficiency by increasing temperature [11] is prevented. For the experiments, the LED pairing B and I is used with a lighting time of six minutes and a post-curing time of 20 minutes. The operating current is varied in a range from 0.06 A to 1 A.

For the experiments, the LED pairing B and I is used with a lighting time of six minutes and a post-curing time of 20 minutes. The operating current is varied in a range from 0.06 A to 1 A. The resulting graph is shown in Figure 6. Here, no significant differences for the surface hardness are shown in a corridor of the operating current IB between 0.25 A and 1 A. The surface hardness decreases strongly below an operating current of 0.25 A, from 29.5 to 27.3 HV1 within a range of 0.125 A. Than sharply from 27.3 to 20.8 HV1 within a range of 0.06 A.

Otherwise, the experimental conditions, e.g., distance between test body and light guides, are kept constant. Because the experiments for determining a suitable LED pairing at room temperature did not provide a clear statement, they are repeated in parts. For this purpose, all four UV LEDs were combined with one LED each at the lower and upper blue wavelength spectrum, the LEDs E and I. The results of the first series of measurements at a curing temperature TA of 60 ° C, a lighting time tB of 6 min, a post-curing time of 20 min, and an operating current IB of 0.5 A are shown in Figure 7. It draws the conclusion that LED B has the best suitability for the process, with a surface hardness of 64 HV1.

It draws the conclusion that LED B has the best suitability for the process, with a surface hardness of 64 HV1. The desired reference surface hardness of > 60 HV1 is achieved for the first time during the entire experiment procedure.1714HG [HVl) 70, 065, 065,060, 055, 052,050, 045, 04 0, 0w ci5 <(ca<(w ci5 alcaalw ci5 uca uw ci5 CcaCFigure 7: Total hardness HG of eight LED pairings, lighting time tB = 6 min, current IB = 0,5 A,post-curing time th = 20 min, curing temperature TA = 60 ° C, number of LEDs n = 2LED B is chosen for the UV range for further testing. Therefore LED B is combined with all blue LEDs and the experiment is repeated with the results displayed in Figure 8. The graph shows that the results of the pairing B & E (49.2 HV1) and B & I (54.9 HV1) deviate from the results of the previous test series (B & E = 65 HV1; B & I = 64 HV1). This deviation can be attributed in part to a shorter warm-up phase of the test bodies.

This deviation can be attributed in part to a shorter warm-up phase of the test bodies. By reaching a new maximum surface hardness of 66.9 HV1, the LED pairing B & F is selected for further tests.HG [HVl) 70,065, 060, 055, 054,950, 045, 04 0, 0w ci5 alL.I. ci5 alI!) ci5 al::c: ci5 alcaalFigure 8: Total hardness HG of five LED pairings, lighting time tB = 6 min, current IB = 0,5 A,post-curing time th = 20 min, curing temperature TA = 60 ° C, number of LEDs n = 21715In order to validate the results, further test series were carried out which are not considered in detail. As a conclusion of these experiments, it can be stated that, above all, the temperature increase has a decisive influence on the achievement of the surface hardness of the reference process. Therefore, the focus is on the influence of the curing temperature TA in the following step.

Therefore, the focus is on the influence of the curing temperature TA in the following step. The temperature TA will be varied between 21 °C room temperature and 60 °C and a preheating time of tV of 6 min for the test body is set. As shown in Figure 9, the resulting surface hardness HG decreases exponentially as a function of the curing temperature TA from 64.7 HV1 at 60 °C to 27.3 HV1 at 21 °C.

As shown in Figure 9, the resulting surface hardness HG decreases exponentially as a function of the curing temperature TA from 64.7 HV1 at 60 °C to 27.3 HV1 at 21 °C. This underlines the cure temperature TA being the strongest process factor so far.Ho [HVl) 80,070,064,760,050,040,030,027,320,021305060Figure 9: Total hardness HG of the LED B&F as a function of the curing temperature TA, lightingtime tB = 6 min, current IB = 0,5 A, post-curing time th = 20 min, number of LEDs n = 8Measurement of surface hardness over lighting time of preheated sample Having found the appropriate light source constellation LED B & F, examined the influence of radiation intensity IB and the curing temperature TA, and after increasing the number n of LEDs, one more step to identify the possibility of an inline curing process is to minimise the lighting time tB.

In this experiment, the test body is preheated for tV = 6 min and removed from the heat source directly after lighting time. The curing temperature was increased to 70 °C. The resulting graph in Figure 10 shows a lighting time of two seconds and a surface hardness of 46.1 HV1. After 30 seconds, a hardness of 68.5 HV1 is measured and represents a maximum value in these experiments. The value set by the reference process > 60 HV1 is measured after a lighting time of 8 sec. Furthermore, a sample was heated at 70 °C for six minutes without exposure to a lighting process.

The value set by the reference process > 60 HV1 is measured after a lighting time of 8 sec. Furthermore, a sample was heated at 70 °C for six minutes without exposure to a lighting process. Even after a waiting time of 20 minutes after the process, the composite was still pasty.1716HG [HVl] 75, 070, 068,565, 060, 055, 05 0, 04 5, 04 0, 024ts [s)81530Figure 10: Total hardness HG of the LED B&F as a function of the lighting time tB, preheating time tV = 6 min, current IB = 0,5 A, post-curing time th = 20 min, curing temperature 70 °C, number of LEDs n = 8DiscussionThe wavelength measurement results conclude in a potential for optimisation of the curing process in terms of curing time. The limited number of wavelengths considered in LED selection offers room for further optimization and use of more appropriate wavelength LEDs is a first conclusion to be drawn.

The limited number of wavelengths considered in LED selection offers room for further optimization and use of more appropriate wavelength LEDs is a first conclusion to be drawn. In contrast, a reason for the temporal modulation of the radiation intensity remains unclear. It could promote the curing process, regulate the process temperature by regulating the energy input, or reduce the thermal load on the LEDs and the printed circuit board.

It could promote the curing process, regulate the process temperature by regulating the energy input, or reduce the thermal load on the LEDs and the printed circuit board. Following the steps taken to determine the critical factors of the curing process and the additionally investigated factor preheating time tV, the following factors are listed in Table 2.Table 2 : Critical factors for the curing process SymbolCritical FactorRangeUnitWavelength Number of LEDs Operating current Lighting time Curing temperature Preheating timeλUV und λBl N IB tB TA tV385 (B) & 458 (F) 2 – 8 ≥ 0,5 8 – 360 60 – 70 0 - 360[nm] - [A] [s] [°C] [s]The first factor, the wavelengths λUV and λBl respectively the LED configuration of the process, is consistent with the spectral absorption peaks at 380 and 468 nm of the two photo- initiators Lucirin TPO and Kamherchinon used in Enamel.

Individual LED characteristics and manufacturer-dependent quality differences of the luminous efficacy at the same operating currents IB could influence the results. There is no significant difference in the total hardness HG of the1717samples correlating with the number of LEDs n. A negative influence of a higher number of LEDs on the curing process is excluded, but a reduction in the number of light sources could make sense in the context of a cost-effectiveness analysis of the overall process. The operating current IB and the radiation intensity of the LED light sources proved to be sufficient, starting from a value of 0.5 A. A further increase of these could neither accelerate the hardening process nor achieve better hardness values. The determination of a more exact threshold of the operating current may be economically viable for the continuous operation of a 3D printer.

The determination of a more exact threshold of the operating current may be economically viable for the continuous operation of a 3D printer. The lighting time tB can be selected depending on preheating time tV and hardening temperature TA between eight and 360 seconds, with a correlation of these three with each other. Thus, the UV and blue light radiation entry into the material proved sufficient after a lighting time of 8 sec, implied that a curing temperature TA of 70 °C and a sufficiently high preheating time tV is selected. This result suggests, that the 12 min lighting and curing time of the reference process is decreased by almost 99% down to 8 sec with comparable surface hardness value. It is safe to say that the reference process incudes a safety factor to ensure high reliability. A safety factor has not been included in the new process. The preheating time tV is also a limiting component.

A safety factor has not been included in the new process. The preheating time tV is also a limiting component. If the sample to be cured is not heated sufficiently, the curing process and thus the lighting time tB take longer, which results in a longer heat input. Since, during these experiments, the highest examined curing temperature was 70 °C, it is unclear how the curing process reacts to a further increase in temperature. However, according to the product data sheet of the manufacturer Vita, it should not exceed a temperature of 120 °C. Other problems in the process handling may occur if the curing time is too short and the curing process is thereby too fast. For example, the deposition process of the material to the surface be negatively influenced at the nozzle where the material is exposed to the light for the first time. The material behavior during a dynamic movement of an extrusion based AM process within a machinery needs to be examined.

The material behavior during a dynamic movement of an extrusion based AM process within a machinery needs to be examined. The influence of the deposition strategy considering parameters such as nozzle diameter, travel speed and volume flow will increase the process complexity. The transferability of the results with standard test specimens to individually deposited strands must be evaluated and can not be considered as given.Summary and outlookIn this work, the key process parameters for an inline curing process within a 3D printing process are determined. An approved reference process is examined and a hardware setup is developed, which offers the possibility to be integrated into machinery. The dependencies of various process parameters among each other could be specified qualitatively. In particular, the curing temperature, the lighting time, and a possible preheating time of the body to be cured play significant roles in the process.

In particular, the curing temperature, the lighting time, and a possible preheating time of the body to be cured play significant roles in the process. The overall hardness, as the average surface hardness of the top and bottom of the specimen, is suitable as a quality standard for the evaluation of curing processes. The lighting and curing time 12 min of the reference process is decreased by almost 99% down to 8 sec with comparable surface hardness value.Sustainable potential for further optimisation is provided by the three parameters illumination time, curing temperature, and sample preheating time. They all have significant effects on the curing process and influence each other during the process. With the need of further examinations, it is still unclear which technology can optimally ensure a constant curing temperature in this AM process or how a specific preheating can be implemented and what impact this has on the body.

Additionally, defects and deformations may develop with a time delay to the laser energy input. Thus, currently, the component quality is only determinable after the finished process.Here, active laser thermography, a nondestructive testing method, is adapted to PBF-LB/M, using the defocused process laser as heat source. The testing can be performed layer by layer throughout the manufacturing process. We study our proposed testing method along experiments carried out on a custom research PBF-LB/M machine using infrared (IR) cameras.Our work enables a shift from post-process testing of components towards in-situ testing during the AM process. The actual component quality is evaluated in the process chamber and defects can be detected between layers.Introduction and MotivationAdditive Manufacturing (AM) of metals as a novel and complex technique is still prone to defects and unwanted process deviations causing them.

Defects in the finished components may include cracks [1], porosity [2], and geometric deviations, among others. These defects can lead to reduced component properties and quality [3], so that time-consuming reworks and even rejects become unavoidable. Therefore, a suited quality control for AM is necessary. This is currently done in two ways: on the one hand, via in-situ monitoring of the AM process itself, e.g., of the melt pool over time. This is admittedly helpful as process quality and component quality are indeed directly related, but not sufficient as the actual component status during manufacturing is not fully determinable. On the other hand, currently, a subsequent ex-situ testing is necessary to determine the actual component status, e.g., via X-ray computer tomography.

On the other hand, currently, a subsequent ex-situ testing is necessary to determine the actual component status, e.g., via X-ray computer tomography. For safety critical components this testing is needed, but is also very time-consuming, and rejects are detected only after the finished AM process.Due to this, a high optimization potential lies within shifting from ex-situ testing towards nondestructive in-situ testing (NDT). The actual quality and status of the built AM component would become determinable during the process. For this, we introduce the possibility of using active laser thermography (or active laser thermographic testing, TT) as an NDT procedure in each layer of the build process. This novel approach is applied on laser powder bed fusion of metals (PBF-LB/M) with 316L stainless steel in the following. For comparison purposes, the common in-situ monitoring technique of optical tomography (OT) is also used simultaneously in each layer.

For comparison purposes, the common in-situ monitoring technique of optical tomography (OT) is also used simultaneously in each layer. In the following, after further general insights and the experimental methods, the results of OT as monitoring and TT as testing are compared and discussed. Thereby, an initial assessment of the in- situ TT for AM approach is possible, and gathered insights may be used for future works.1978BackgroundIn the following, the two used techniques for AM process and component assessment are introduced: optical tomography as process monitoring method on the one hand, and active laser thermographic testing as nondestructive testing method on the other hand.Optical Tomography (OT)OT is a current AM process monitoring technique used for PBF-LB/M. [4] Classically, a single camera in the visual or near-infrared wavelength range is monitoring the powder bed. It is used for long-time exposures throughout a complete manufactured layer.

[4] Classically, a single camera in the visual or near-infrared wavelength range is monitoring the powder bed. It is used for long-time exposures throughout a complete manufactured layer. In this way, the emitted radiation of the AM process in the observed wavelength range is captured. Due to the integration over time, no temporal information is available, but highly spatially resolved data for the whole layer. From this, conclusions about the local process quality can be drawn. It currently finds rise in popularity due to low hardware costs as well as a simple general usage.Active Laser Thermographic Testing (TT)For the NDT technique of active laser thermography, heat is introduced via a laser onto the surface of the tested object. The resulting thermal response of the unsteady heat flux is captured via an IR camera. Using additional algorithms for analysis, defects are detectable in this way due to several occurring effects.

The resulting thermal response of the unsteady heat flux is captured via an IR camera. Using additional algorithms for analysis, defects are detectable in this way due to several occurring effects. These effects include altered local thermal properties or altered local laser absorptivity, among others. The approach of scanning over the test object’s surface with a laser spot is often referred to as flying spot thermography. With it, surface defects like open cracks are detectable. [5] This approach was for example also applied on the PBF-LB/M process by Herzer and Schilp utilizing the process laser. [6] However, the introduced energy per area was rather high (no NDT approach), and solely information of a single testing scan track and not the scanning sequence as a whole was taken into account. In a prior work, we could show on a test specimen that nondestructive testing inside a PBF-LB/M machine is possible via defocusing the process laser.

In a prior work, we could show on a test specimen that nondestructive testing inside a PBF-LB/M machine is possible via defocusing the process laser. [7] Based on this, we incorporate this approach into the actual AM process, and take another step towards a nondestructive in-situ testing for PBF-LB/M.Experimental Setup and MethodsThis section describes the setup and methods used in this study: the PBF-LB/M research machine, the camera setups (OT and TT), the AM build job along with the incorporated artificial defects, and the thermographic testing routine. Furthermore, the analysis of the recorded data is described.PBF-LB/M Research Machine - SAMMIETo optimize controllability and data accessibility, a noncommercial PBF-LB/M research machine was successfully developed at the BAM facilities in Berlin-Adlershof. The Sensor-based Additive Manufacturing Machine (SAMMIE) was set up in this study as illustrated in Figure 1.

The Sensor-based Additive Manufacturing Machine (SAMMIE) was set up in this study as illustrated in Figure 1. It portrays the elements of SAMMIE, beginning with the laser path including focus unit (varioSCAN by SCANLAB) and a galvanometer scanner (intelliSCAN by SCANLAB). The laser source employed is an ytterbium continuous wave fiber laser (IPG Photonics) operating at 1070 nm,1979capable of generating a maximum nominal power of 500 W. Control over the manufacturing process and the laser path components is achieved through the Direct Machining Control (DMC) software. A shielding gas system with filter ensures a process at low oxygen levels. Finally, the configuration includes a process chamber with a powder bed unit inside with a build volume of 50 x 70 x 35 mm³. An outstanding feature of SAMMIE is its capacity to host multiple camera systems simultaneously during the active process.

An outstanding feature of SAMMIE is its capacity to host multiple camera systems simultaneously during the active process. The off-axis arrangement of the cameras at the process chamber’s optical entry points allows observing the entire powder bed from a close proximity. Furthermore, it is feasible to implement on-axis camera installations along the laser path, enabling precise monitoring of the laser affected zone. However, although not utilized in this publication, the on-axis capability is intended for future investigations. The entire PBF-LB/M process operation is controlled by LabVIEW-based software (National Instruments) which ensures a fully accessible AM process with SAMMIE.

The entire PBF-LB/M process operation is controlled by LabVIEW-based software (National Instruments) which ensures a fully accessible AM process with SAMMIE. Moreover, additional procedures may be incorporated such as active laser thermographic testing after a finished layer.Figure 1: Overview of the utilized PBF-LB/M research machine SAMMIE.Camera SetupsFor the comparison of monitoring and testing, an optical tomography and a thermographic testing setup were installed at SAMMIE. The OT setup comprised of a ZWO ASI 6200 MM Pro Mono camera with a full-frame resolution of 62 Mpx (9576 x 6388 px²) at 16-bit image depth. The camera was mounted on top of the process chamber using a 100 mm objective (cf. Figure 1). This resulted in a spatial resolution of approximately 10 µm/px when focusing on the build platform. Filters were installed in the optical path to ensure optimal conditions for the recorded images. Firstly, a bandpass filter at 725 nm was included.

Filters were installed in the optical path to ensure optimal conditions for the recorded images. Firstly, a bandpass filter at 725 nm was included. Secondly, neutral density (ND) filters at varying strengths were incorporated via a filter wheel to switch between them effortlessly. The ND filter suited best was a ND2.2 (0.63% transmittance) at 725 nm. With this setup, monochromatic OT longtime exposures at full dynamic range were possible.1980For the thermographic testing data acquisition, the Goldeye CL-033 TEC1 camera by Allied Vision operating in the short-wave infrared (SWIR) range was used (cf. also Figure 1). Its built-in InGaAs sensor gives a spectral sensitivity spanning from 0.9 µm to 1.7 µm. The camera captures images at a full-frame resolution of 640 x 512 px² with a 14-bit image depth and a frame rate of 234 Hz. For focusing on the build plane, the camera was equipped with a 100 mm SWIR lens.

For focusing on the build plane, the camera was equipped with a 100 mm SWIR lens. An extension tube, approximately 35 mm in length, was included to achieve the desired focus. Consequently, a spatial resolution of 40 µm/px was achieved. Moreover, a notch filter (NF1064-44 by Thorlabs) at the laser wavelength was positioned in front of the camera objective. This ensured that only the thermal response was captured while excluding the reflected laser itself. The camera system was installed off-axis at one of the optical access points of the SAMMIE process chamber using an adjustable mirror system. This configuration facilitated alignment between the recorded camera images and the region of interest on the build plane.Build Job and DefectsThe PBF-LB/M build process was performed on the research machine SAMMIE with 316L stainless steel powder (1.4404) as base material. Particle sizes ranged from 10 µm up to 45 µm (d50 = 32.55 µm).

Particle sizes ranged from 10 µm up to 45 µm (d50 = 32.55 µm). The process itself was conducted under argon shielding gas atmosphere at a laser power of 167 W and a scanning speed of 700 mm/s (monodirectional) while the layer height was set to 40 µm. The hatching distance between scan tracks was at 100 µm while the laser spot diameter (D86) was at approximately 80 µm.Throughout the manufacturing process of the monitored and tested component, two separate zones with artificial defects were implemented. These defects were omitted areas where no scanning and therefore no powder melting took place. The first zone comprised 24 round defects (cf. Figure 2a) with 12 cylindrical (top of the figure) and 12 spherical shapes (bottom of the figure) which ranged from layers 75 to 87. In this zone, laser scanning directions changed by 67 degrees between each layer. Defect diameters were at 1.5 mm down to 400 µm in steps of 100 µm.

In this zone, laser scanning directions changed by 67 degrees between each layer. Defect diameters were at 1.5 mm down to 400 µm in steps of 100 µm. The second zone included defective areas in the form of narrow triangles and notches (cf. Figure 2b). This artificially recreates (powder filled) crack structures in the AM component. They were included in the layers 140 to 149 and are open to the top. The laser scanning directions were kept constant between layers and parallel to the crack structures to ensure precise manufacturing. To even enhance the right edge of the left triangle, the edge was scanned separately with 3 degrees scanning angle. The two triangles had a maximum width of 1.1 mm. The notch widths ranged from 400 µm to 100 µm in steps of 100 µm while their length was at 8 mm.

The two triangles had a maximum width of 1.1 mm. The notch widths ranged from 400 µm to 100 µm in steps of 100 µm while their length was at 8 mm. With these two defect zones, different types of defects were analyzable.Figure 2: Internal artificial defects in the manufactured component with cylindrical and spherical defects in the first zone (a) and notch defects in the second zone (b).1981Thermographic Testing and AnalysisWhile the OT solely required a longtime exposure during manufacturing of a layer, the active laser thermography required an additional energy input via the defocused process laser. This scanning over the component’s surface was executed after every layer in an area of 8.5 x 16.5 mm² right across the defective regions. Appropriate scanning parameters had to be found empirically to ensure an energy input as low as possible for the sake of non-destructiveness of testing.

Appropriate scanning parameters had to be found empirically to ensure an energy input as low as possible for the sake of non-destructiveness of testing. However, energy input had to be high enough that the SWIR camera was able to capture the thermal response. The utilized thermographic testing parameters were at 32 W laser power with a scanning speed of 22 mm/s (monodirectional). The distance between scan tracks was 250 µm which resulted in 67 lines for the tested area. The defocused process laser had a spot diameter (D86) of approximately 1.1 mm. The off-axis SWIR camera was recording with a framerate of 235 Hz at an exposure time of texp = 200 µs with an aperture of 22.For the analysis of the thermographic data, the maximum thermograms were investigated. For every pixel, the maximum thermal response in the whole recorded testing sequence was found and displayed to create the maximum thermogram. This was done for every layer.

For every pixel, the maximum thermal response in the whole recorded testing sequence was found and displayed to create the maximum thermogram. This was done for every layer. Background is that the information content of a single frame is rather low, so the sequence as a whole is taken into account. The highest recorded thermal response for a specific pixel (which corresponds to a specific location on the AM component surface) gives information on the regional thermal properties. With this, surface defects can be localized as they alter the thermal properties of the AM component locally. Additionally, to assess the potential of detecting subsurface defects by active laser thermography, the sum thermogram was investigated. In this case, the sum of the recorded thermal signal was calculated for every pixel. This also takes the cooling behavior into account. It was investigated for layer 89, so two layers after the round defects were closed by sound, complete layers.

This also takes the cooling behavior into account. It was investigated for layer 89, so two layers after the round defects were closed by sound, complete layers. Therefore, the nominal defect depth was at 80 µm. To enhance contrast in the sum thermogram, the exposure time of the SWIR camera was increased to texp = 500 µs for this layer. As both, optical tomography and thermographic testing, were conducted off-axis from different positions, perspective transformations were performed. [8] In this way, optical distortions were removed and comparability between monitoring and testing was enhanced.Results and DiscussionIn the following, the obtained images of the optical tomography as monitoring method and the thermographic testing as nondestructive testing approach are compared. This is done layer by layer for the first zone with round defects (layers 86 to 89) and for the second zone with notch defects (layer 144).

This is done layer by layer for the first zone with round defects (layers 86 to 89) and for the second zone with notch defects (layer 144). Goal is to determine the current capabilities of the in-situ TT as a novel approach for PBF-LB/M in comparison to the OT as in-situ monitoring.Round DefectsFigure 3 shows the obtained images for layer 86 of the long-time exposure OT on the left (a) and the maximum thermogram of the TT on the right (b). The perspective transformation corrected perspective distortions introduced by the off-axis configurations for both methods. This allows an improved comparison. In the OT image, missing process scan tracks are apparent. This was due to the capture delay of the used high-resolution OT camera, thus scans at the beginning of the layer manufacturing are missing. However, the comparison of OT and TT was still feasible, and1982therefore, the data recording was deemed successful.

However, the comparison of OT and TT was still feasible, and1982therefore, the data recording was deemed successful. The OT clearly shows the omitted areas of the artificial round defects. All diameters down to 400 µm were resolved with ease due to the high spatial resolution of the utilized camera.Figure 3: Optical Tomography (a) and Thermographic Testing (b) results of layer 86.Cylindrical and spherical artificial defects are at full diameters.The incorporation of an in-situ thermographic testing routine into the PBF-LB/M process was successful as Figure 3b shows. This marks a great step on the way to include in-situ NDT into AM because to our knowledge, a nondestructive TT implementation without remelting was not achieved before. The selection of the SWIR spectral range for the IR camera was motivated by the planned on-axis experiments including active laser thermography.

The selection of the SWIR spectral range for the IR camera was motivated by the planned on-axis experiments including active laser thermography. The laser path, specifically the focus unit, is optimized for transmissivity at a laser wavelength of 1070 nm which makes the neighboring SWIR range most suitable for a transfer to on-axis operation. However, in the current investigation, off-axis experiments were conducted to explore the general feasibility of the approach. Furthermore, the same SWIR camera setup would allow for additional in-process monitoring of the melt pool. Additionally, SWIR cameras are comparatively more affordable than cooled mid-wave infrared cameras (MWIR).The maximum thermogram in Figure 3b depicts pixel-wise the maximally emitted infrared radiation for the finished layer during the scanning of the active laser thermography with flying spot method. For this, the color scale marks lower to higher recorded values with blue to yellow.

For this, the color scale marks lower to higher recorded values with blue to yellow. It shows a thermal signature for all of the round defects in this layer. This thermal response was caused by the interaction between the defocused process laser of the active laser thermography scanning and the loose, unmelted powder. In these areas, the maximum of the emitted infrared radiation is higher compared to the sound, solidified areas of the AM component. With this, even the smallest defects at a diameter of 400 µm are visible. Therefore, OT as in-situ monitoring and TT as in-situ NDT showed very well aligning results for defect detection. However, the thermal signature was distinctively weak for the smallest defects which might have two possible explanations: either the spatial resolution of the SWIR camera was at its limit or the actual diameter of the artificial defect in the finished layer was smaller than the nominal 400 µm.

The latter might be caused by the small, omitted defect areas melting together which changes the defect appearance and reduces its diameter. However, an actual representation of the finished component with its1983defect signatures (e.g., via X-ray computer tomography) would be needed. The maximum thermogram furthermore indicates on its right side the edge between component and powder bed. The defocused process laser for the scanning routine of the active laser thermography slightly struck the powder bed here. This resulted in a visible thermal response. This might be utilizable in the future to determine the actual positioning of the manufactured component in the powder bed and in the process chamber.Figure 4: Optical Tomography (a) and Thermographic Testing (b) results of layer 87.Cylindrical artificial defects are at full diameters, spherical defects begin closing.Figure 4 shows the results for OT and TT of the subsequent layer 87.

On the left (a), the OT exhibits a very similar outcome to the layer before. The missing scan tracks shifted in their position due to the hatching angle changing between layers by 67 degrees. The omitted artificial defect areas are all fully visible once more. Even the top of the spherical round defects with diameters well below 400 µm are visible. Again, the high resolution of the utilized OT camera plays a part in this. The TT results on the right (b) are similar to the layer before, but also show differences. The largest of the cylindrical defects are visible in the maximum thermogram. However, the thermal signatures of the smaller defects blend in with a decisive thermal response of the layer’s surface. Due to this, the smallest defects are not clearly visible in the thermogram. Especially the top of the spherical defects with their diameters below 400 µm are not identifiable with the overlaying response pattern.

Especially the top of the spherical defects with their diameters below 400 µm are not identifiable with the overlaying response pattern. Where this response stemmed from is not completely clear, but an explanation could be the rough nature of the AM surface. As the flying spot thermography is a surface NDT technique, surface inhomogeneities can negatively affect the capabilities of finding structural defects like cracks; and as severe roughness is prevalent in PBF-LB/M, the occurring thermal response as influencing factor must be under control for in-situ TT in the future. Additionally, the effect of small artificial defects melting together plays once more a role here as described above. The defect’s actual extent is not clear solely from the OT as monitoring, and a ground truth would be needed.1984Figure 5: Optical Tomography (a) and Thermographic Testing (b) results of layer 88.

Firstcompletely manufactured layer after the incorporated artificial round defects.Layer 88 is the first complete layer without artificial defects after the defect zone. With it, the capabilities of the two applied methods (OT and TT) to find subsurface defects (in this case at 40 µm layer height) was evaluated; Figure 5 shows the results. The OT on the left (a) indicates slight signatures where the largest defects were located in the layer before. So, the emitted radiation during the build process was altered by the underlying defects and was captured integratively by the OT camera. The signatures are visible with the human eye, but they are not very distinct. This might open the possibility for future usage of artificial intelligence (AI) algorithms to detect subsurface defect signatures in OT - e.g., via convolutional neural networks (CNN).

This might open the possibility for future usage of artificial intelligence (AI) algorithms to detect subsurface defect signatures in OT - e.g., via convolutional neural networks (CNN). So, there is potential in high-resolution OT for quality assurance in future works.In contrast, the maximum thermogram of layer 88 in Figure 5b does not show any signs of the subsurface defects, indicating that the chosen setting was not capable of detecting them. This is not unexpected as flying spot thermography is an active laser thermography technique that is specifically used to detect surface defects like open cracks. However, other laser excitation strategies beyond simple monodirectional scanning motions might be evaluated in the future for in-situ NDT in AM. These active laser thermography strategies are more suitable for finding subsurface defects. [9] Interestingly enough, the maximum thermogram showed a rather homogeneous thermal response across the component’s surface.

[9] Interestingly enough, the maximum thermogram showed a rather homogeneous thermal response across the component’s surface. This contrasts the distinct thermal pattern in the subjacent layer 87 (Figure 4b). It therefore aligns more with the homogeneous thermal response of layer 86 (Figure 3b). The reason for this deviation across layers is not known at this point.1985Figure 6: Optical Tomography (a) and Thermographic Testing (sum thermogram, b) results oflayer 89. Second complete cover layer after the incorporated artificial round defects.Finally, for cylindrical and spherical defects, layer 89 was analyzed - so two layers above the artificial defects at a layer height of 40 µm per layer. The results are shown in Figure 6. The OT on the left (a) still contains defect signatures of the largest artificial round defects. Obviously, they are less pronounced than before with another layer in between.

The OT on the left (a) still contains defect signatures of the largest artificial round defects. Obviously, they are less pronounced than before with another layer in between. However, it shows that a subsurface defect detection via monochromatic high-resolution OT might hold potential even across multiple layers.For the TT, Figure 6b depicts the sum thermogram of layer 89. This is contrary to the maximum thermograms of Figure 3b to Figure 5b before. The exposure time texp of the SWIR camera was increased from 200 µs to 500 µs for the sum thermogram. This higher texp would result in an overexposed maximum thermogram, but for the sum thermogram, the sensitivity is beneficially increased. When considering Figure 6b, the sum thermogram shows slight defect signatures for the largest of the round defects. The results and capabilities are therefore comparable to the monitoring via OT.

When considering Figure 6b, the sum thermogram shows slight defect signatures for the largest of the round defects. The results and capabilities are therefore comparable to the monitoring via OT. This indicates that a nondestructive in-situ detection of subsurface defects via active laser thermography during the AM process is possible. Background is that the sum thermogram holds information about the cooling process after the energy input of the active laser thermography. This improves the capabilities of detecting subsurface defects as defects alter the local thermal properties (e.g., cooling rate). This is then indicated in the sum thermogram. The approach holds therefore similarities to the integrative method of the OT. However, it again must be distinguished between OT as monitoring and TT as testing. The former records information about the process while the latter nondestructively assesses the actual final component.

The former records information about the process while the latter nondestructively assesses the actual final component. As mentioned before, the laser excitation strategy of TT can be adapted to further improve finding subsurface defects. Furthermore, comparisons of TT results and X-ray computer tomography are crucial to further evaluate the capabilities of in-situ TT. Both aspects are planned for future works.Notch DefectsThe second defect zone from layer 140 to 149 incorporated the artificial triangle and notch defects. Exemplarily, the results for layer 144 are depicted in Figure 7. The OT image on the left (a) is missing the first scan tracks because of the capture delay of the high-resolution OT camera. This1986is identical to the OT images discussed beforehand for the round defects. Again, an additional delay would have prevented this and will be included for future works.

This1986is identical to the OT images discussed beforehand for the round defects. Again, an additional delay would have prevented this and will be included for future works. Due to this, the left triangle defect structure was not recorded since its right edge was scanned separately before the rest of the layer at a different scanning angle. However, the rest of the image was fully captured and a comparison to TT is possible. The image shows that the right triangle as well as all four notches were resolved by the OT. This also includes the narrowest notch at 100 µm width. This shows that the OT as process monitoring was able to capture also fine elongated artificial defects in this configuration.Figure 7: Optical Tomography (a) and Thermographic Testing (b) results of layer 144. Artificialtriangle and notch defects are fully visible.Figure 7b shows the maximum thermogram of the active laser thermography similarly to layers 86 to 88.

Artificialtriangle and notch defects are fully visible.Figure 7b shows the maximum thermogram of the active laser thermography similarly to layers 86 to 88. The border of the powder bed and therefore the edge of the component is visible once more, most distinctly on the right edge. Both triangle defects are visible in full. Between them, an additional thermal response is located. The origin of this pattern is not identifiable at this point. All four notches were detected, even the narrowest at 100 µm. For the two notches at the lowest widths (100 µm and 200 µm), the thermal response in the maximum thermogram distinctively dropped compared to the other two, slightly wider notches. This might indicate that an effect of melting together occurred for these fine structures. Nevertheless, the TT was also able to detect all of the elongated artificial defects.

This might indicate that an effect of melting together occurred for these fine structures. Nevertheless, the TT was also able to detect all of the elongated artificial defects. In general, it showed once more that the TT as testing approach was capable of achieving comparable results in defect detection to the OT as process monitoring procedure.The received layer information from the maximum thermograms can be stacked to obtain a quasi-3D data representation of the manufactured component. A visual representation of this is depicted in Figure 8a for the layers 140 to 149. A comparison to the 3D information from computer tomography would be feasible with this. Additionally, Figure 8b shows a photograph of the finished component. On the top, the artificial defects with triangles and notches of the second defect zone are visible. Furthermore, it shows the powder stuck in the elongated defects.

On the top, the artificial defects with triangles and notches of the second defect zone are visible. Furthermore, it shows the powder stuck in the elongated defects. It was sintered by the laser scanning of the TT and a clear increase in surface roughness is visible. Above the upper component surface, a cover layer of sintered powder was easily removed which was attached to the component. The TT was continued automatically with the PBF-LB/M process for several layers even though the investigated component was finished. This created the cover of sintered powder1987and increased the surface roughness of the finished component. This further underlines the need to examine the influence of the TT’s heat input on the final component quality.

This further underlines the need to examine the influence of the TT’s heat input on the final component quality. This examination along with the reduction of the mentioned heat input will be studied in the future.Figure 8: Layer-wise quasi-3D information from the TT for the notch defect zone at layers 140to 149 (a) and finished component after the process (b).Summary and OutlookThis work investigated the novel approach of including active laser thermography (TT) as nondestructive testing method into the AM process of PBF-LB/M. For this, the defocused process laser at low laser power scanned over the surface of a finished layer. The thermal response was captured via a short-wave infrared camera to detect defective regions. The capabilities of this in-situ TT approach were assessed by simultaneous comparison with optical tomography (OT). This in-situ monitoring method observes the AM process via a long-time exposure for every layer.

This in-situ monitoring method observes the AM process via a long-time exposure for every layer. In contrast, the TT is testing the component after each finished layer. For the investigations, a square component with artificial defects was manufactured via PBF-LB/M. The defects comprised omitted areas with round and elongated (notch) shapes at various sizes. The component was manufactured with the research machine SAMMIE using 316L stainless steel powder. The inclusion of the TT sequence after each layer was fully successful.The results of OT and TT regarding defect display aligned very well. The two systems could successfully determine open artificial defects in AM layers. This was the case for both, round and notch defects. Even notches with a nominal width of 100 µm were detectable.

The two systems could successfully determine open artificial defects in AM layers. This was the case for both, round and notch defects. Even notches with a nominal width of 100 µm were detectable. For finding subsurface defects, TT with sum thermograms instead of maximum thermograms showed more promising results, but further investigation is needed.The true extent of the defects after the process was not determinable here, and a closing of the defects by extended melting might have occurred. Due to this, a core aspect of future works on the topic will be the referencing to X-ray computer tomography (XCT) as a ground truth. The testing results from TT and the actual component information from XCT will be compared directly. Furthermore, the influence of the additional heat input on the AM component will be investigated. A reduction might be achievable by optimizing TT parameters while using cameras sensitive at longer wavelengths.

A reduction might be achievable by optimizing TT parameters while using cameras sensitive at longer wavelengths. The latter would decrease the minimum surface temperature at which the infrared camera captures a thermal response. Additionally, using an on-axis instead of an off-axis configuration will be beneficial for transferring the in-situ TT approach onto commercial1988PBF-LB/M machines and for improved spatial resolutions. Finally, the capabilities of detecting subsurface defects might be improved by adapting the laser excitation strategy to a more sophisticated and well-suited method. Consequently, a big variety of investigations and improvements are still needed.

Consequently, a big variety of investigations and improvements are still needed. However, the results presented within this work mark an essential step on the road from time-consuming and expensive ex-situ testing towards a nondestructive in-situ testing for AM processes.AcknowledgementThe project ATLAMP is supported by the Federal Ministry for Economic Affairs and ClimateAction (BMWK) on the basis of a decision by the German Bundestag.Grant numbers - BAM: 16KN086124; THETASCAN GmbH: 16KN086125.References[1] G. Mohr, J. Johannsen, D. Knoop, E. Gärtner, K. Hummert and C. Emmelmann, "Processing of a high-strength Al-Fe-Ni alloy using laser beam melting and its potential for in-situ graded mechanical properties," Proceedings of the Lasers in Manufacturing Conference, 2017. [2] S. Oster, P. P. Breese, A. Ulbricht, G. Mohr and S. J.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceIn-situ Detection of Build Defects in Additive Manufacturing via Impedance-BasedMonitoringLogan Sturm1, Mohammed Albakri2, Dr. Christopher B. Williams1, Dr. Pablo Tarazaga2 1 Design, Research, and Education for Additive Manufacturing Systems Lab 2 Vibrations, Adaptive Structures and Testing Laboratory Department of Mechanical Engineering, Virginia Polytechnic and State UniversityAbstractIn this paper, the authors explore the use of impedance-based monitoring techniques for the in- situ detection of additive manufacturing build defects. By physically coupling a piezoceramic (PZT) sensor to the part being fabricated, the measured electrical impedance of the PZT can be directly linked to the mechanical impedance of the part.

By physically coupling a piezoceramic (PZT) sensor to the part being fabricated, the measured electrical impedance of the PZT can be directly linked to the mechanical impedance of the part. It is hypothesized that one can detect in- situ defects of part mass and stiffness by comparing the signatures collected during printing of parts with that of a defect-free control sample. In this paper, the authors explore the layer-to- layer sensitivity of this technique. A control sample is created using Material Jetting and the change in signatures between various layer intervals is measured. To evaluate the technique’s ability to perform in-situ detection, several parts containing designed defects (e.g., internal voids) are fabricated and their layer-to-layer signatures are compared to a control sample. Using this technique, the authors demonstrate an ability to track print progress and detect defects as they occur.1.

Using this technique, the authors demonstrate an ability to track print progress and detect defects as they occur.1. IntroductionWith the development and maturation of additive manufacturing (AM) there are an increasing number of end use products being fabricated with these technologies. This shift from producing prototypes to fabricating functional end-use products has created a growing need for part validation and quality control [1]. Defects or process variations that might be acceptable in a demonstrational, tooling, or prototype part are no longer acceptable in an end use product. In particular, the aerospace industry has begun to adopt AM technologies for the fabrication of high value, low volume parts.

In particular, the aerospace industry has begun to adopt AM technologies for the fabrication of high value, low volume parts. Some example of AM parts being produced and tested include direct printed metals (e.g., Rolls Royce Trent-XWB bearing (Figure 1a) and RUAG’s Sentinel satellite bracket (Figure 1b)) and polymers (e.g., FAA-approved ULTEM 9085 aircraft air duct created by Stratasys and Orbis (Figure 1c)).1458(a) Rolls Royce Trent- XWB bearing [2](b) RUAG’sbracket [3]Sentinelsatellite(c) Stratasys/Orbis’ conformalair duct [4]Figure 1. Example applications of AM to fabricate mission-critical aerospace components.1.2 In-situ Monitoring for Additive ManufacturingIn their Measurement Science Roadmap for Metal-based Additive Manufacturing, NIST states that “Existing NDE techniques are not optimized for AM processes, materials or parts. Techniques are lacking for in-situ NDI, and post-process AM part inspection.” [1].

Techniques are lacking for in-situ NDI, and post-process AM part inspection.” [1]. To address the needs presented by NIST, the authors suggest that an ideal in-situ NDE technique for AM would have the following attributes:Cost-effective.  Able to be conducted quickly.  Able to evaluate parts irrespective of geometry and material. In other words, the technique would be able to analyze inaccessible features, large parts, fine features, rough surface finish, etc.Able to detect both surface and internal defects. In particular, the ability to detect changes in non-surface material properties that may occur due to factors such as thermal effects during or after the fabrication of subsequent layers.In addition to this, the authors propose that an ideal method would be resistant to cyber-attacks. As discussed in their previous work [5], there are variety of cyber-physical vulnerabilities that exist in quality control (QC) and AM due to the digital nature of the process chain.

As discussed in their previous work [5], there are variety of cyber-physical vulnerabilities that exist in quality control (QC) and AM due to the digital nature of the process chain. Since a Stuxnet-style attack that compromises a printer could compromise the sensors on the printer as well, it is advantageous to have independent side-channel measurements that can monitor the part fabrication independently from the printer.Current in-situ monitoring approaches have focused on the performance of direct metal systems such as SLM and LENS. A variety of approaches have be presented in research such as high- speed cameras [6], thermocouples, pyrometers [7, 8]. These techniques are limited to the1459inspection of the surface geometry of a part. Other methods of inspection have been considered, such neutron diffraction and X-ray backscatter [9]. The methods can detect defects volumetrically, but require expensive radiation sources.

Other methods of inspection have been considered, such neutron diffraction and X-ray backscatter [9]. The methods can detect defects volumetrically, but require expensive radiation sources. Laser ultrasonic testing has also been investigated as a way of detecting surface defects and has some capability to detect volumetric properties such as thickness [10]. Many in-situ systems focus on monitoring the melt pool instead of the entire build area since this allows for increased resolution. One method presented by Lott et al. is through the use of an illumination laser in combination with a high-speed camera [11].

One method presented by Lott et al. is through the use of an illumination laser in combination with a high-speed camera [11]. While this allows for precise monitoring of the melt pool it is not able to monitor the entire build area or to detect thermal changes that may occur in previous layers as a result of the heating of already fabricated layers during the fabrication of new layers.Some work outside of direct metal AM has been performed using a variety of sensors and the use of data fusion to integrate their results together. Rao et al. demonstrated the use of thermocouples, accelerometers, an IR temperature sensor, and a borescope to monitor and detect build defects in an FDM process [12]. The primary focus of in-situ monitoring so far has been on monitoring either the build layer or the machine behavior. In the first case, the primary concern is with tracking the geometry of the part or with monitoring the thermal characteristics to model the behavior of the system.

In the first case, the primary concern is with tracking the geometry of the part or with monitoring the thermal characteristics to model the behavior of the system. While these methods are useful, they lack the ability to directly measure the material properties and are limited to monitoring the surface of the part. In the second case the focus is on detecting when the machine is operating outside of normal parameters (indicating a build failure). While this method provides valuable feedback information, it does nothing to directly monitor the properties of the part.1.3 Research GoalBased on the previous work that has been done in in-situ monitoring, the authors believe that there is still a need for an in-situ monitoring technique that is able to directly interrogate the material properties and geometry of the part being fabricated.

The overall goal of this paper is assess the feasibility of using in-situ electromechanical impedance measurements for NDE of additively manufactured parts. The authors’ previous work has established that piezo-ceramics sensors can be used as a post-process NDE tool to detect defects in fully fabricated AM parts [13]. Based on this work, the authors believe that impedance based measurements may have potential for in-situ monitoring as well. The use of these sensors in-situ poses new challenges. One concern is that, since the part is attached to a build plate during fabrication, the impedance response due to the much larger mass of the machine may overwhelm the relatively small mass of the fabricated part, which could result in defects being undetectable. Another concern is that vibrations from the machine might influence the sensor reading, resulting in a poor signal to noise ratio.

Another concern is that vibrations from the machine might influence the sensor reading, resulting in a poor signal to noise ratio. Finally, there is the concern that the location of a part on the build tray might affect the measurement, and thus prevent successful comparison of the measurements between different parts. To evaluate if impedance-based measurements via piezo-ceramics are a valid method for in-situ AM measurement, the authors sought to verify the following hypotheses:14601. For a single part at a given layer, the sensors will be able to repeatedly measure the same signature.2. For a single part, the sensors will be able to detect the change from one layer to another. This change will be greater than the signature variation in a single layer.3. While being fabricated, all parts will have similar signatures until the defects occur.

This change will be greater than the signature variation in a single layer.3. While being fabricated, all parts will have similar signatures until the defects occur. After the defects are introduced, the sensors will be able to detect differences in the parts’ signatures.To verify these hypotheses, the authors designed and printed a series of test specimens which contain internal build errors (as described in Section 3.1). Impedance measurements were conducted and compared against a set of defect-free parts (Section 3.2). Results from this comparison are presented and analyzed in Section 4. Closure and future work are presented in Section 5.The need for reliable methods of verifying the quality of AM parts presents a significant barrier to widespread industrial adoption of AM technologies. Before AM parts can be adopted in end use and mission critical roles reliable methods of certification are needed.

Before AM parts can be adopted in end use and mission critical roles reliable methods of certification are needed. While some existing methods such as X-ray CT have been able to provide validation for existing parts, these methods are expensive and new designs are pushing the size limits of current scanners. Several organizations such as NIST [1], NASA [14], and the ASTM F-42 committee [15] have recognized the need for NDE techniques optimized for AM.To address this research gap, the authors propose the use of electromechanically impedance measurements as a means to detect and identify AM defects. The authors’ previous work has demonstrated this potential with fully fabricated part and the authors seek to extend this work to in-situ monitoring. By monitoring parts in-situ defects are able to be detected sooner allowing for reduced waste and potentially the repair of defects before they become internal to the part.

By monitoring parts in-situ defects are able to be detected sooner allowing for reduced waste and potentially the repair of defects before they become internal to the part. A detailed overview of impedance-based SHM and its extension to NDE of AM parts is provided in Section 2.2.0 Non-Destructive Evaluation for Additive Manufacturing2.1 Electromechanical Impedance for SHM and NDEThe basis for using electromechanical impedance for structural health monitoring or non- destructive evaluation is the impedance signature that a part will have. Each part has a unique combination of mass, stiffness, and damping characteristics that result in a unique dynamic response (impedance signature). If damage occurs to the part or a defect is introduced these part characteristics will change, altering the dynamic response. In impedance based SHM has shown potential as a real time damage assessment method that is non-intrusive, cost effective, and highly sensitive [16].

In impedance based SHM has shown potential as a real time damage assessment method that is non-intrusive, cost effective, and highly sensitive [16]. This technique utilizes pieozoelectirc materials, specifically lead zirconate titanate (PZT) wafers as both the sensor and the actuator to both excite the structure and measure its response [17,18]. By attaching these sensors to the part, the mechanical properties of the part1461are coupled with the electrical properties of the sensor, allowing the measurement of the mechanical impedance of the part to be simplified to the measurement of the electrical impedance of the sensor which is directly related. The electrical impedance of the piezo can be easily measured, allowing changes in the mechanical impedance to be detected.

The electrical impedance of the piezo can be easily measured, allowing changes in the mechanical impedance to be detected. This allows changes in the part characteristics due to factors such as build defects to be detected by changes in the electrical impedance signature [17, 18].Figure 2 shows a schematic of a piezoelectric patch attached to a printed part, which is represented by an equivalent spring, mass, and damper system. Assuming linear piezoelectricity, the constitutive equations of the piezoelectric materials operating in 1-3 mode can be expressed as [19]Figure 2. A piezoelectric patch attached to a mechanical structure represented by a spring- mass-damper system.𝐸 𝜎11 + 𝑑13𝐸3𝜀11 = 𝑠11Eq.

A piezoelectric patch attached to a mechanical structure represented by a spring- mass-damper system.𝐸 𝜎11 + 𝑑13𝐸3𝜀11 = 𝑠11Eq. 1𝐷3 = (𝑑𝑇)31𝜎11 + 𝜖33𝜎 𝐸3 ,where 𝜀11 is the strain tensor component in the 1-direction, 𝜎11 is the corresponding component of the work-conjugate stress tensor, 𝐷3 is the electric displacement in the 3-direction, 𝐸3 is the 𝐸 is the complex electric field in the 3-direction, 𝑑13 is the piezoelectric coupling coefficient, 𝑠11 mechanical compliance of the material measured at zero electric field. For the simplified 1D, 𝐸 reduces to the inverse of Young’s modulus of elasticity of the plane-stress problem at hand, 𝑠11 piezoelectric material.

For the simplified 1D, 𝐸 reduces to the inverse of Young’s modulus of elasticity of the plane-stress problem at hand, 𝑠11 piezoelectric material. 𝜖33𝜎 is the complex permittivity measured at zero stress.Due to the coupled electromechanical behavior of piezoelectric materials, the electrical impedance of the piezoelectric sensors is directly related to the mechanical impedance of the host structure, 𝑍𝑠𝑡 , as follows [9]𝑍(𝜔) = [𝑖𝜔𝑏𝑙 ℎ2 𝑑13 𝐸 ( ( 𝑠11𝑡𝑎𝑛(𝑘𝑙) 𝑘𝑙(𝑍𝑝𝑧𝑡 𝑍𝑝𝑧𝑡 + 𝑍𝑠𝑡𝜎 )]) − 1) + 𝜖̅33−1,Eq. 21462where 𝑍𝑝𝑧𝑡 = −𝑖𝑏𝑙ℎ (𝑠11𝐸 𝜔𝑡𝑎𝑛(𝑘𝑙)𝑘𝑙)−1is the piezoelectric sensor short circuit impedance, where𝐸 is the wave number, 𝜌 is the density of the piezoelectric material, 𝑏, ℎ and 2𝑙 are𝑘 = 𝜔√𝜌𝑠̅11 the piezoelectric patch width, thickness and length, respectively.Because this method is based on vibration the sensitivity of the measurements is dependent on the frequency range used.

For a defect to be detected the wavelength of the excitation signal needs to be smaller than the characteristic length of the defect [19]. To achieve higher sensitivity using the impedance based method higher frequency excitations should be used. In work done by Peairs et al. the use of preselected frequency ranges based on the free response of the piezoelectric sensors [20]. Based on this work, they concluded that both the characteristics of the piezoelectric sensor and the part are important for determining the optimal frequency range.The use of electromechanical impedance measurements has been demonstrated in a variety of structural applications. Park et al. used this technique for applications such as bridge sections, pipe joints, and reinforced concreate walls [21].

Park et al. used this technique for applications such as bridge sections, pipe joints, and reinforced concreate walls [21]. Other studies have used impedance measurements to detect defects in both real world conditions and laboratory environments [22 23].The authors hypothesize that electromechanically impedance measurements may be particularly well suited for in-situ monitoring of AM parts. The authors’ previous work has demonstrated the potential for this technique to be used on fully fabricated parts [SFF piezo]. The use of the technique in-situ shares many of the same benefits such as being cost effect (i.e no radiation source needed), time efficient (i.e. each sweep takes seconds to run), and the ability to detect changes in material properties. The features of impedance based measurements allow it to fit many of the needs identified in section 1.2. By using high frequencies, small defects can be detected layer-by-layer as the part is fabricated.

By using high frequencies, small defects can be detected layer-by-layer as the part is fabricated. Further, because the method is able to detect changes throughout the volume of the part, material changes that might occur in already fabricated layers can be detected.3. MethodTo utilize impedance based techniques for in-situ inspection of AM parts the authors propose a system where an initial set of control parts is used to set a baseline for detecting defects. In this system several copies of the designed part are printed and monitored using the impedance based techniques. For each layer where inspection is desired the signature is recorded. These control parts are then tested using nondestructive and destructive techniques to ensure that they were properly fabricated and meet all design requirements. The signatures from these validated parts are then used to create a set of baseline signatures.

The signatures from these validated parts are then used to create a set of baseline signatures. Subsequent parts are monitored using theses impedance based techniques and are compared to the baseline signatures. Any defects that occur will alter the signature resulting in a difference from the baseline that can be detected. For this method to be applicable the hypotheses presented in Section 1 need to be validated.14633.1 Test specimen Design and FabricationAs previously presented in Section 1 this study sought to validate three base hypotheses:1. For a single part at a given layer, the sensors will be able to repeatedly measure the same signature.2. For a single part, the sensors will be able to detect the change from one layer to another. This change will be greater than the signature variation in a single layer.3. While being fabricated, all parts will have similar signatures until the defects occur.

This change will be greater than the signature variation in a single layer.3. While being fabricated, all parts will have similar signatures until the defects occur. After the defects are introduced, the sensors will be able to detect differences in the parts’ signatures.To investigate these hypotheses, the authors designed three test parts, (i) a control (Figure 3a, “C”), (ii) a large rectangular prism cavity (Figure 3b, “S”), and (iii) a triangular prism cavity (Figure 3c, “R”). Each part also contained a small cavity at its base for embedding the piezo- ceramic sensor/actuator. The test parts have overall dimensions of 35mm x 19mm x 15mm. The thickness of the wall containing the defects is 1mm. The “S” defect is 33mm x 17mm x 9.5mm for a total volume of 5330mm3 (53.4% of total volume) and a per layer volume of 16.8mm3 (0.17% of total volume). The “R” has a right triangular cross section with sides 33mm and 9.5mm and a depth of 17mm.

The “R” has a right triangular cross section with sides 33mm and 9.5mm and a depth of 17mm. The total volume of the defect is 2665mm3 (26.7% of total volume) and the smallest layer change is 0.027 mm3 (0.00027% of total volume). The piezos are embedded after 2.5mm of the parts have been printed and the defects begin after 4.5mm.ABCFigure 3. Test parts A) Control sample “C”, B) Rectangular prism cavity “S”, C) Triangular prism cavity “R”To minimize the potential variation from naturally occurring defects in the printing process as well as from part damage during embedding the design of the test parts was kept as straightforward as possible. This allowed for simple visual inspection to be performed during fabrication to ensure that the parts were being fabricated as expected. The small lower cavities are open and are designed for support material to be removed to allow the piezos to be embedded in the parts.

The small lower cavities are open and are designed for support material to be removed to allow the piezos to be embedded in the parts. The defect cavities are enclosed to allow the parts to be cleaned after fabrication without removing the support material contained inside the cavities. This allows the total mass of1464the part to remain constant while the difference in composition between model and support material changes. The large rectangular cavity is designed to cause a defect with a constantly increasing size, while the triangular cavity will cause a linearly increasing void size. The goal of this linearly increasing defect is to allow for a more precise determination of the minimum detectable defect size.The parts were printed on a Connex350 printer using VeroWhite+ material and a matte finish. This process was chosen due to the ability to easily mount and access the sensors.

This process was chosen due to the ability to easily mount and access the sensors. The process also has a high resolution, 300 x 600 DPI and 0.03mm layers, which allows for measuring the effect of smaller amounts of material being added. The support material also has similar density to the model material meaning that the effect of a stiffness change without a mass change can be investigated. The primary drawback of the Polyjet process is the limited material selection. The model material is a photopolymer that is less rigid than materials available on other processes, in particular those on metal based systems. Since the transmission of vibration through a part is directly linked to the stiffness of the material this lack of rigidity has the potential to reduce the effective range and sensitivity of the sensors.

The support material is very soft and has the potential to act as a large dampener, absorbing the vibrations generated by the piezo-ceramics and further reducing the sensitivity of the sensors. Despite these material limitations the Polyjet process provides a good testbed since the move to more ridged materials would be expected to improve the performance of the sensors in many cases. The test parts were arrayed in a three-by- three grid on the build tray, with one of each type of sample design in every row and column, as shown in Figure 4. The parts were printed until 2.5mm , at which the print was paused, support material was removed from the cavities, and the piezo-ceramic sensors were placed inside and bonded to the part with an adhesive. Before embedding, copper tape was attached to the piezos to provide connection points for measurements.To minimize variation between sample measurements, each piezo used was cut from the same sheet.

Leads were soldered to the ends of the copper tape to provide reliable attachment points. The print was then resumed, with measurements being taken at the layers as shown in Table 1. To allow measurements to be taken consistently the authors paused the build at each listed layer before taking measurements. This was to avoid any changes that could be caused by the print head depositing more material while a measurement was being taken. While the pause and measurement operation was performed manually in this study, it could be automatically programmed into the system to provide interference free in-situ monitoring. In addition to the measurements taken in-situ, the final parts were measured on the build tray, after being removed from the build tray (but before having excess support material removed) and after having been cleaned from support material.1465C1S1R1R2C2S2S3R3C3Figure 4. Test parts with piezos embedded and layout of the test parts.

Test parts with piezos embedded and layout of the test parts. The letter indicates the type of part (e.g. “C” for control, “S” for the rectangular prism, and “R” for the triangular prism. The number indicates the sample number for each part. In this case three samples of each type were printed.Table 1. Part layers where measurements were taken. For each layer, the volume of model material in each part is shown.

The right column shows the percentage difference between the defect parts and the control at each layer, indicating the relative size of the defectVolume mm3Volume %LayerCSRSR150229422942294100.0%100.0%160249424942494100.0%100.0%19832523033324793.3%99.9%20834513064343788.8%99.6%21836513095362284.8%99.2%22838503127380181.2%98.7%23840503158397578.0%98.2%24842493189414475.0%97.5%26846483251446569.9%96.1%28850473314476665.7%94.4%31856463407517660.4%91.7%34862443501553956.1%88.7%39872423657603750.5%83.4%44882393813640346.3%77.7%49892373969663643.0%71.8%14663.2 Analysis techniquesIn this study, the authors situated the analysis in a supervised learning context, where differences in the impedance signatures of fabricated parts against a baseline signature from a known, defect free part are compared.

Two of the control samples were used to establish the baseline. The third control sample was used to validate this baseline.To validate the hypotheses put forth in section 1 three steps were needed, (i) to determine the repeatability of measurements for a single part on a single layer, (ii) to determine the change in a single part between layers, (iii) to determine the difference between parts on the same layer.Determining the repeatability of the measurements of a single part on the same layer establishes a baseline noise threshold. This baseline also indicates whether or not there is significant interference from the machine between measurements since the authors’ previous work has shown good repeatability on fully fabricated parts. To achieve this, three measurements were taken of each part at each layer and the signatures were compared at each frequency to obtain maximum and average deviations for each part.

To achieve this, three measurements were taken of each part at each layer and the signatures were compared at each frequency to obtain maximum and average deviations for each part. After this the three measurements are added together and all impedance signatures are shifted vertically such that their average value (excluding impedance peaks), match to compensate for any inconsistency in the connectors’ resistivity (e.g., due to variations in soldering and wire length).To quantify the difference in signatures between layers, a damage metric defined as𝑅𝑀𝑆𝐷 = √∑(𝑍𝐷−𝑍𝐵𝐿)2 2 𝑍𝐵𝐿Where ZD is the real component of the impedance signature of the part being tested, and ZBL is the real component of the baseline impedance signature, and n is the total number of data points in the impedance signature was used. This metric was used in two ways, (i) to compare a single part to itself in subsequent layers and (ii) to compare different parts to each other in the same layer.

This metric was used in two ways, (i) to compare a single part to itself in subsequent layers and (ii) to compare different parts to each other in the same layer. In the first case, the signature of the first measured layer is used as the baseline and each subsequent layer is compared to this first layer to determine an overall change in the part. This metric is also used to compare each measured layer to the layer immediately preceding it to determine the change between layers. In the second case, the signatures of the first two control samples are averaged together and used as the baseline. The third control sample is used to verify the baseline value and the other samples are all compared to this baseline. Since each sample is designed to be identical before the defect is introduced the initial baseline difference can be used to calibrate for natural variation between the piezo.

Since each sample is designed to be identical before the defect is introduced the initial baseline difference can be used to calibrate for natural variation between the piezo. This was to account for any difference that might exist due to factors such as differences in the piezoceramics, mounting, or build tray location.1467(2)4. Results4.1 Evaluating RepeatabilityIn order to perform in-situ measurements it was necessary to determine if the machine would cause interference between measurements and what the background noise level for variation was. As stated in hypothesis 1, it was necessary to test if the machine was causing the measurement of a single part and layer to vary from one moment to another. Figure 5a shows a typical example of a set of three measurements taken from layer 150 of a single control sample, in this case C1. The average deviation of this set of measurements was 0.167Ω, indicating strong repeatability.

The average deviation of this set of measurements was 0.167Ω, indicating strong repeatability. Based on this, minimal machine interference can be expected between measurements, while some bias may occur, it will be consistent across measurements. The highest deviation was found to be 7.32Ω in layer 448 of part R3. As shown in Figure 5b, this is mostly the result of a linear offset occurring between measurements. A linear offset will be caused by an increase in the total resistance of the system usually attributable to something such as an increased connection resistance in the case of a poor wire connection. This offset can be corrected for by subtracting the mean of each trace to normalize the data. Overall, the samples showed good repeatability of measurements throughout all of the layers tested. Of note are the higher deviations for R1 and R3, indicating potential challenges in comparing these measurements.

Of note are the higher deviations for R1 and R3, indicating potential challenges in comparing these measurements. This is most likely due to errors that occurred while embedding and is discussed in more detail in Section 4.3.Figure 5. Part signatures sample parts. On the left is a typical measurement deviation shown for Control 1 (maximum deviation = 0.49Ω). On the right is an example of the worst case deviation observed, in this case for Ramp 3 (maximum deviation = 7.3Ω).1468Table 2. Repeatability measurements for each sample taken across all layers measured.

Repeatability measurements for each sample taken across all layers measured. Of noteare the particularly high numbers for R1 and R3.Part C1 C2 C3 R1 R2 R3 S1 S2 S3Max Deviation (Ω) 2.524 1.224 0.891 3.862 0.735 7.323 1.211 0.824 1.241Average Deviation (Ω) 0.271 0.089 0.054 0.855 0.140 1.210 0.111 0.091 0.0554.2 Evaluating Change in a Single Part Between LayersTo evaluate the proposed technique’s ability to detect changes between layers in a single part, as proposed in the second hypothesis (Section 1), the authors compared both (i) the total change from the first layer to the measured layer, and (ii) the change between layers for a single sample. This experimental approach provided a measure of total change and of relative change between measured layers.To compare the total change the authors used a single part (C1) and selected the signature of the first measured layer (layer 150). This layer was set as the reference layer to which all other layers were compared.

This layer was set as the reference layer to which all other layers were compared. This layer was 3.42mm into the part and was chosen to allow the piezo to be completely enclosed and secured by model material. This layer occurs before any defects are introduced into the part. When comparing the total change in signature from the reference layer, it is observed that all layers showed a significant difference, far above the noise threshold. This indicates that changes could easily be detected. Figure X shows the total difference from the reference layer for part C1. Increasing difference in signatures can be seen as more layers are added to the part. The smallest difference that was measured was between 10 layers (0.3mm) with the added volume being 199.5mm3. A sample of the measured signatures is shown in Figure X. A slight shift (left) can be seen in the peaks of the signatures as more layers are added to the part.1469Figure 6.

A sample of the measured signatures is shown in Figure X. A slight shift (left) can be seen in the peaks of the signatures as more layers are added to the part.1469Figure 6. Total difference in signatures of part C1 from each measured layer to the referencelayer (150). Increasing difference can be seen as more material is added to the part.Figure 7. Select signatures of part C1. Layers 150, 160, 228, and 498 are shown. A slight shift(left) of the peaks can be observed with increasing layers.1470It is important to note that while two layers may have a similar total magnitude of difference from the reference layer, their signatures might be significantly different from each other. To illustrate this, imagine a sine wave as the reference layer.

To illustrate this, imagine a sine wave as the reference layer. While both a leftward and a rightward shift of the signature by five degrees causes the same magnitude of change, the two signatures are indicating very different changes, and the difference between the two shifted signatures is greater than the difference from the reference part. To evaluate this, the signature of each layer in C1 was compared to the signature of the layer before it. Figure 8 shows the difference of each layer from the layer before it; in each case the difference is well above the noise threshold. Because the separation of measured layers is not constant throughout the print, it is important to look at the effects of volume of material added on the measured difference. In a direct comparison between the addition of 10 and 20 layers the difference from adding 20 layers will be inflated due to more material being added.

In a direct comparison between the addition of 10 and 20 layers the difference from adding 20 layers will be inflated due to more material being added. Since each layer has a cross section that contains the same area of material, two different measurements can be equally compared by normalizing to the number of layers added.Figure 9 shows the difference in signatures for C1 between layers divided by the number of layers added. In this normalized graph, a trend toward decreasing sensitivity can be seen to start occurring around layers 268 to 288. At the earlier layers the signatures tend to change more significantly, as more material is added there tends to be less movement and variation in the signatures. This explains the large variation and unpredictability in the difference of the earlier layers compared to the more stable trend of the later layers. This trend can probably be attributed to the increased amount of total mass in the part.

This trend can probably be attributed to the increased amount of total mass in the part. As more material is added the volume/mass of each added layer remains constant, while the total mass of the part continues to increase. This causes the percent change of the part to decrease with each layer. In addition the new material is further away from the sensor. As distance from the sensor increases the ability to detect changes will decrease. .The results of this test show that the difference between layers in a single part can be detected and that this difference is above the variation threshold found when verifying the first hypothesis. This method is able to successfully detect increasing change from a reference layer and is also able to clearly detect change between layers as proposed in the second hypothesis. The diminishing layer to layer difference indicates that the hypothesis is valid only with a certain range of part masses.

The diminishing layer to layer difference indicates that the hypothesis is valid only with a certain range of part masses. A sufficiently large system my result in a sensitivity drop below the minimum detection threshold.1471Figure 8. Difference in signature of each measured layer to the measured layer before it.Figure 9. Difference in signature after scaling for the mass difference being measured. After the initial variation it can be seen that a trend toward decreasing sensitivity with added part mass begins to occur.14724.3 Defect Detection Comparison Across PartsTo determine if defects could be detected compared to a reference “good” part, a baseline measurement was created by averaging the measurements of two of the control samples, C1 and C2 (Section 3.2). Due to small variations in the piezos and in their mounting some difference exists between the baseline and the test parts before the defects are introduced.

Due to small variations in the piezos and in their mounting some difference exists between the baseline and the test parts before the defects are introduced. To calibrate the measurements, the difference found in the first measurement is subtracted from the subsequent measurements. This was done to minimize any bias due to the sensor. Figure 10 is a color graph showing the difference magnitude (RMSD) for each part across all layers. Darker blues indicate less difference between the part and the baseline. Bright yellows indicate increased difference from the baseline.It can be seen that the control samples maintain a relatively small difference from the baseline while the samples with the defects have a small initial difference, but quickly show a much larger difference from the baseline. The red line indicates where the defect occurs and the circled area shows where the defect can be detected.

The red line indicates where the defect occurs and the circled area shows where the defect can be detected. For the rectangular prism void the defect can be detected at layer 198 (defect volume 219 mm3, corresponding to a 6.7% volume fraction of the total part). As the defect continues to grow, it can be seen that the difference between the defect parts and the control parts continues to trend towards increasing, with the correlation between similar parts being fairly close.Figure 10. Difference between the baseline (average of C1 and C2), the control samples, the triangular prism void (R), and the rectangular prism void (S) for each measured layer. The samples with defects can be seen to increase in difference while the controls remain approximately the same. The defect in the S samples can be detected in layer 198, the first measured layer after the defect is introduce at layer 186.

The defect in the S samples can be detected in layer 198, the first measured layer after the defect is introduce at layer 186. The R defect can be detected in layer 218.1473In looking at the initial difference for the ramp samples, it was found that R1 and R3 had a very high initial difference. Upon investigating the signatures, shown in Figure 10, it is clear that there is a significant difference in these samples compared to the control. Due to the important of the physical bond between the part and the sensor, the potential for damage to the piezo, and the difficulty in completely removing support material and attaching the sensor it is likely that this error occurred as part of the embedding process. Due to the apparent defect in embedding, these samples were excluded and only R2 is shown for the difference comparison is subsequent layers.(a)(b)Figure 10. a) Comparison of the measured signals of the controls samples. Signatures are similar and closely clustered.

a) Comparison of the measured signals of the controls samples. Signatures are similar and closely clustered. b) Comparison of the measured signals of the triangular prism samples. The R1 and R3 samples show a significant difference from the expected signature. This is likely due to a failure in the embedding process.The remaining sample (R2), as expected, showed a smaller degree of difference than the previous rectangular void samples (S1,S2,S3), due to the smaller size of the defect. In the triangular prism sample, the defect could be accurately detected in layer 218, corresponding to a defect assize of 29mm3 (0.8% volume fraction of the total part). Before this, the difference was not significant enough to determine with certainty that a defect had occurred. After detection, and with increasing defect size, the detected RMSD difference generally continued to increase, but to a lesser level than the rectangular prism samples.

After detection, and with increasing defect size, the detected RMSD difference generally continued to increase, but to a lesser level than the rectangular prism samples. Since the triangular prism defect is smaller this result is expected. Figure 12 shows the color chart comparison for the ramp for each layer.1474Figure 12. Difference between the baseline, the control samples, and the support samples for each measured layer. The support samples can be seen to increase in difference while the controls remain approximately the same. The defect can be detected in layer 218.5.0 Closure and future Work The authors have shown that piezo-ceramic sensors can be used in-situ to monitor parts. In this study three hypotheses has been verified, the repeatability of the same measurement, the detection of changes in a single part between layers, and the detection of differences between parts after a defect is introduced.

For the first hypothesis, within the same part and layer, the results were very repeatable with the average deviation of the control sample’s damage metric being only 0.27Ω (Table 2). This showed that there was minimal interference from the machine between measurements, a requirement for being able to perform high sensitivity in-situ measurements.For the second hypothesis, the method was able to detect changes occurring between one layer to another. The detected resolution was ten layers (0.3mm), corresponding to a volume change of 199.5 mm3 (4.2% volume fraction). Much smaller changes should be detectable, however this was the smallest increment that was tested in this study. In general, the difference from the base layer tended to increase as more layers were added to the part. Sensitivity of the measurements was also noted to decrease as more material was added to the part.

Sensitivity of the measurements was also noted to decrease as more material was added to the part. While these trends are not conclusive these results reinforce initial hypothesis on the behavior of the system.For the third hypothesis, the sensors proved to be less effective at detecting defects between parts in-situ than in similar tests on completed parts. The smallest defect that was detected in-situ was 29mm3 (0.8% volume fraction), compared to defects as small as 8mm3 in fully fabricated parts. The authors hypothesis that this is due in part to the extra supporting material around the part, the extra mass in the system due to being fixed to the build tray, and the potential for the location on the build tray to have an effect on the signature. Despite the somewhat reduced sensitivity, the1475in-situ method was still able to reliably detect defects in the parts as they occurred.

Despite the somewhat reduced sensitivity, the1475in-situ method was still able to reliably detect defects in the parts as they occurred. Refinements to the technique and analysis should be able to allow for the detection of smaller defects.The parts in this study were fabricated on a Connex 350. While this system provides a convenient test bed for validating the concept of impedance based in-situ monitoring, the materials available on this system are expected to perform poorly. Due to the high damping of the support material on this system it is expected that he sensitivity of the method could be improved. Detection of defects is dependent of the transmission of vibration through the part; therefore parts made out of stiffer materials can be expected to perform better with this method.

Detection of defects is dependent of the transmission of vibration through the part; therefore parts made out of stiffer materials can be expected to perform better with this method. In particular, the authors hypothesize that metal based AM systems will perform significantly better and will allow for the detection of smaller defects.To more fully qualify the sensitivity and size of defects that can be detected in-situ further work needs to be done investigating smaller step sizes. In addition, the authors plan to conduct in-situ measurements on an FDM system with the hypothesis that the stiffer support material may provide less dampening and result in more sensitive detection abilities. Other future work includes the validation of these methods on metal systems where the much greater stiffness is expected to improve the sensitivity of the method. To further refine the sensitivity of the method more sophisticated analysis techniques than RMSD are being investigated.

One example is the use of peak detection and comparison to more precisely isolate the important sections of the signature.Acknowledgements:This material is based upon work supported by the national science foundation under grant #1446804: “CPS Synergy: Collaborative Research: Cyber-Physical Approaches to Advanced Manufacturing Security” Any opinions, findings, and conclusions or recommendations expressed in this material are those of the author(s) and do not necessarily reflect the views of the National Science Foundation.1476References [1] NIST, 2013, “Measurement Science Roadmap for Metal-based Additive Manufacturing,” byhttp://www.nist.gov/el/isd/upload/NISTAdd_Mfg_Report_FINAL-2.pdf, Energetics Incorporated, May.prepared[2] Rolls Royce, 2015, “High-Powered Trent XWB-97,” Rolls Royce Media Insightshttp://www.rolls-royce.com/media/insights/simon-burr.aspx, August 31.[3] RUAG, 2014, “RUAG pioneers 3D printing for space,” RUAG Media Release http://www.ruag.com/group/media/media-releases/news/ruag-pioneers-3d-printing-for- space/d39b1226395457f3362a92216064fa1e/ August 31.[4] Stratasys, 2015, “Additive Manufacturing for Aerospace: FAA-Approved Air Duct for ‘Flying Eye Hospital’ Produced in Just Days,” http://blog.stratasys.com/2015/03/05/3d- printed-air-duct-flying-eye-hospital/ March 5.[5] A.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceIn-situ Electrical Resistance Measurements for Soldering Studies in Hybrid AMAlexander P. Pustinger, Joselin Corral, Arianna Villegas, David EspalinThe University of Texas at El Paso, W.M. Keck Center for 3D Innovation, El Paso, TexasKeywords: Hybrid AM, 4-wire resistance measurements, encapsulation, laser solderingThis report was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights.

Reference herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof: The views and opinions of-authors expressed herein do not necessarily state or reflect those of the-United States Government or any agency thereof.AbstractThe convergence of additive manufacturing (AM) along with multiple applicable technologies has been shown to augment the functionality of printed parts such that mechanical, electrical, and electromagnetic functions can, for example, reside within the same part. This work used an electrical connector commonly used in satellites to demonstrate a hybrid AM approach by embedding the connector, ultrasonically embedding wires, and laser soldering. The connector was a six-pin Easy-on FFC/FPC connector with a fine pitch and leads of 200-micron width.

The connector was a six-pin Easy-on FFC/FPC connector with a fine pitch and leads of 200-micron width. Resistance measurements were carried out during laser soldering, application of insulation material, and encapsulation with thermoplastic extruded at high temperatures (> 300⁰C). The measurements were made to observe the impact of 1) a laser power profile and 2) thermal input from the extrusion nozzle and heated build environment for diagnostics of low yield or performance. Additionally, temperature history data was collected to coincide with resistance measurements.IntroductionThe interest in hybrid additive manufacturing (hybrid AM) for on-orbit production has grown due to its potential to fabricate a wide range of devices, such as small satellites and space station electronics, using a single manufacturing system. This approach offers the advantage of reducing the need for on-orbit spare part inventory.

This approach offers the advantage of reducing the need for on-orbit spare part inventory. Of particular significance is the fabrication of electronics on orbit, which is relevant in both terrestrial and on-orbit scenarios. Examples of on- orbit fabrication exist both with and without the utilization of AM. For instance, COSM Advanced Manufacturing Systems LLC made significant progress in improving an electron gun, optics, and beam control for in-space assembly using the Electron Beam Free Form Fabrication method. This technique employs E-beams to additively manufacture metal parts for the aerospace industry [Bowman, 2018]. Another example involves the use of deployable structures and on-orbit direct write printing to enhance Supervised Autonomous Assembly (SAA), an approach that promotes autonomous assembly and efficiency in AM without requiring a system operator [Grande, 2021].

In contrast, a non-AM example demonstrates the use of pre-manufactured components assembled1136for on-orbit manufacturing, supporting various operations involved in the design of spacecraft assembly on orbit [Doris, 2018].Despite the notable achievements in on-orbit fabrication, the research focus on printing electronics on-orbit has been limited. To the best of the author’s knowledge, automated hybrid AM has not been utilized for on-orbit electronics fabrication. Therefore, this study aims to fabricate coupons containing components and materials relevant to space applications. These coupons will serve as a baseline for future comparisons with on-orbit or similar environment experiments. Although the fabrication process did not take place on-orbit or in space-like conditions, it considered the potential challenges posed by gravity and vacuum.

Although the fabrication process did not take place on-orbit or in space-like conditions, it considered the potential challenges posed by gravity and vacuum. Before exploring these challenges, a brief overview of 3D electronics and laser soldering will be provided.The utilization of interconnects and conductive materials in AM has progressed to simplify assembly complexity and reduce redundant mass/packaging size. A collaborative effort between NASA/GRC (National Aeronautics and Space Administration/Glenn Research Center) and America Makes resulted in the development of an AM process than embeds interconnects within a hybrid AM process, combining vat photopolymerization (VPP) and direct printing [Liu, 2019]. In their approach, a pressure sensor was fabricated, incorporating VPP-fabricated insulated structures and interconnects created through direct dispensing. By depositing a layer of photocurable material on top, they achieved a fully encapsulated circuit.

By depositing a layer of photocurable material on top, they achieved a fully encapsulated circuit. Other studies have explored the integration of interconnects into flexible materials to create stretchable electronics for applications in soft robotics [Ladd, 2016]. This involved using a syringe extrusion-based 3D printing process with conductive silicone to manufacture a functional game dice with an embedded accelerometer. Additionally, the Aerosol Jet Printing Method was employed in another example, where silver-based ink was deposited onto a PLA prosthetic to create strain gauge sensors [Lu, 2018]. In addition to interconnects, various literature sources have described the use of connectors within AM parts, including SMA connectors [Shemelya, 2015], SMA, coaxial SMA, and N-type connectors [Hao, Yu, Liang, Shemelya, 2018], N-type Coaxial and N-type connectors [Liang, Yu, Shemelya, MacDonald, Xin, 2015], and microUSB connectors [Shemelya, 2015].

Alongside interconnectors and connectors, the quality of the joints between them is crucial for the proper functioning of circuits.To implement laser soldering technology effectively, it is crucial to have a comprehensive understanding of different aspects, such as the resulting microstructure and the interaction between the laser and the solder material [Jiang, 2021]. Qualitative research has revealed distinct intermetallic compound formations (IMC formations) in laser soldering, which is a contactless method that eliminates the need for hot soldering tips. The characteristics of IMC formation are influenced by various factors, including laser wavelength, power output, pulse duration, and laser beam size [Jiang, 2021]. The formation of IMCs significantly impacts the mechanical properties and quality of solder joints. Two types of IMC formations have been observed: scalloped and prismatic.

The formation of IMCs significantly impacts the mechanical properties and quality of solder joints. Two types of IMC formations have been observed: scalloped and prismatic. Scalloped microstructures exhibit refined particles, while prismatic microstructures appear coarse. Solder joints with prismatic IMC formations exhibit higher shear force compared to those with scalloped formations [Jiang, 2021]. The determination of laser parameters that result in either scalloped or prismatic microstructure was achieved using Jackson’s parameter [Jiang, 2021].Numerical analysis conducted by Kunwar (2017) focused on SAC solder paste and compared it to polished microstructure under varying output power and short pulse durations. The1137study utilized a high-power laser with a wavelength of 915 nm. The findings indicated that prismatic IMC formations could be produced by moving the laser at low scan speeds along solder substrates with a power output of 30-50 W (wavelength of 1528-1565 nm).

The findings indicated that prismatic IMC formations could be produced by moving the laser at low scan speeds along solder substrates with a power output of 30-50 W (wavelength of 1528-1565 nm). This dataset was utilized alongside a Neural Network to predict prismatic formations using Jackson’s Parameter [Kunwar, 2020].In-situ monitoring during the laser soldering process has been achieved through X-ray imaging, allowing researchers to observe the morphological changes over time [Ma, 2013]. Comparative studies conducted by Tatsumi (2022) compared the impact properties of solder and microstructure using an infrared wavelength laser and a blue light laser.

Comparative studies conducted by Tatsumi (2022) compared the impact properties of solder and microstructure using an infrared wavelength laser and a blue light laser. The study varied power and pulse durations on a fixed-size joint and found that the efficiency and impact properties were significantly better under blue light due to higher absorption (23% increase compared to IR) [Tatsumi, 2022].A notable aspect to consider regarding laser-material interaction is the monitoring of electrical resistance throughout various stages of joint creation, subsequent process steps involving thermal changes, and the overall lifecycle of the joints. To evaluate the reliability of solder joints under thermo-mechanical cycling, four-wire resistance measurements have been conducted to assess high resistances [Duan, 2014].

To evaluate the reliability of solder joints under thermo-mechanical cycling, four-wire resistance measurements have been conducted to assess high resistances [Duan, 2014]. This study demonstrates a hybrid AM approach that combines material extrusion AM, ultrasonic wire embedding, micro-dispensing of solder paste, and laser soldering to create electrical joints. As part of the process development, in-process resistance measurements were performed across the fabricated joints too characterize their behavior and guide further improvements. This approach was developed to identify the specific process step responsible for joint failure when issues such as high resistance or lack of continuity were observed.

This approach was developed to identify the specific process step responsible for joint failure when issues such as high resistance or lack of continuity were observed. The focus was particularly on understanding the thermal history and resistance response during thermal events in the fabrication process, such as laser soldering and encapsulation with heated thermoplastic material.Materials and MethodologyCoupon FabricationTesting coupons were produced using the Foundry Multi3D System, as described in previous work [Billah, 2019]. The system comprised wo Fortus 400mc material extrusion AM machines from Stratasys, Inc. (Eden Prairie, MN, USA) and a modified LC 3024 Computer Numeric Control (CNC) system from Techno CNC (Ronkonkoma, NY, USA, equipped to support machining, ultrasonic wire embedding, solder paste dispensing, and laser soldering.

The fabrication process followed the following sequence: 1) A polycarbonate (PC) base was 3D printed, 2) a cavity for the connector and a channel for wires were machined into the base, 3) copper wires were ultrasonically embedded into the base, 4) excess wire was machined to the desired length, 5) the connector was placed into the cavity, 6) solder paste was dispensed at the joint, 7) the joint was soldered using laser technology, 8) the area around the connector and atop the PC base was 3D printed to create a cavity, 9) epoxy was dispensed into the created cavity, and 10) the entire PC area, including the epoxy-filled cavity, was 3D printed over.The design of the coupon consisted of a rectangular prism base measuring 35 x 21 x 3.65 mm, featuring a cavity for a 527450697 connector from Molex (Lisle, IL, USA) and channels for bare copper wires with a diameter of 34-gauge (ø = 0.1601 mm), as shown in Figure 1.

The connector used was a six-pin Easy-on FFC/FPC connector with a pitch of 250 microns and leads1138of 200 microns in width. The base was initially 3D printed using PC and then machined to create precise mating features for the connector and wires. An ultrasonic sonotrode was employed to embed the wires into their respective channels. Excess wire measuring 9-14 mm was used to create free leads opposite the connector. The copper wire ends were utilized for resistance measurements, with an expected resistance range of 20.6-25.3 mΩ. During ultrasonic wire embedding, the kinetic energy generated by the sonotrode impaled the wire into the PC material and produced localized heat at the work area. However, the desired level of heating was carefully controlled to ensure that the embedding process occurred in a solid state.

However, the desired level of heating was carefully controlled to ensure that the embedding process occurred in a solid state. This limited heat input and the solid-state nature of ultrasonic wire embedding are considered advantageous for on-orbit manufacturing, as the influence of gravity is minimal and excessive heat can have negative effects.Inside the connector, a separate insertable flexible flat cable of flexible printed circuit (FFC/FPC) was inserted, featuring six leads on the opposite end corresponding to the FFC ribbon. One end of the ribbon was soldered with a 152.4 mm long 24-gauge copper wire to the lead of the connector was measured and found to be approximately 38.6 mΩ. Without solder, the total resistance range was between 59.2-63.9 mΩ. To dispense the solder paste, a Type IV, Sn96.5- Ag3-Cu0.5 solder paste from Nordson (Westlake, OH, USA) was used. The dispensing was carried out by a 7194 Auger style pump, equipped with a 25-gauge needle.

The dispensing was carried out by a 7194 Auger style pump, equipped with a 25-gauge needle. The needle was manually positioned at the appropriate connector lead relative to the tip of the leads, with the assistance of a CMOS camera.Four-Wire Resistance Measurements and Visual ObservationsDuring the process of laser soldering, two CMOS cameras from Basler AG (Ahrensburg, Germany, Europe) and a BK-2841 DC resistance meter from BK Precision (Yorba Linda, CA, USA) were utilized for targeting, solder reflow observations, and four-wire resistance measurements, respectively. To prevent oversaturation from laser light during soldering, the camera used for in-process observations was equipped with a filter. The resistance meter emphasized an accuracy of 0.01% and a resolution of 0.01 μΩ, capable of measuring resistance within the range of 2-200 mΩ (with a maximum output of 5V and 1A at this range).

The resistance meter emphasized an accuracy of 0.01% and a resolution of 0.01 μΩ, capable of measuring resistance within the range of 2-200 mΩ (with a maximum output of 5V and 1A at this range). The resistance meter formed a circuit that incorporated the copper wire, solder joint, connector, and ribbon cable in series, shown in Figure 2. To ensure the proper positioning and stability of the connector, Thermoplastic Polyurethane (TPU) was extruded, acting as a physical constraint in case the cavity did not provide a perfect fit or in the event of movement caused by the resistance meter probes.1139Figure 2. Setup for resistance measurements during laser soldering.Laser SolderingA programmable semiconductor diode laser with a wavelength of 808 nm, manufactured by NaKu Technology, Co., Zhejiang, China, was utilized in conjunction with a LabVIEW virtual instrument (VI) from National Instruments in Austin, TX, USA.

The laser operation followed a specific power profile: 1) Initially, it operated at 2 W until the user corrected the laser position. 2) Then, it gradually ramped up from 2 W to 3.4 W over a span of 15 seconds to induce solder reflow. 3) Subsequently, it ramped down from 3.4 W to 2 W over 15 seconds to allow for solidification. 4) Finally, at the end of the laser soldering process, the laser power was set to 0 W. The laser beam was directed to the tip of the Molex connector lead, with half of the laser spot targeting the copper wire and the other half focused on the lead itself. To ensure synchronization between the laser heating and cooling phases, the resistance meter was triggered to start recording data once the targeting was completed.

To ensure synchronization between the laser heating and cooling phases, the resistance meter was triggered to start recording data once the targeting was completed. Laser soldering is particularly advantageous in on-orbit manufacturing scenarios, as it does not require physical contact, thus reducing the potential impact of gravity effects or extreme reactions to physical forces.EncapsulationFollowing the soldering process, the resistances of all leads were documented at the ambient temperature of the room, which was 23°C. After that, additional PC material was extruded over the embedded wires while forming a cavity measuring 9.97 x 8.125 x 1.27 mm (as shown in Figure 4) to accommodate manually dispensed epoxy material. The top of the coupon was completely covered in a final printing step.

The top of the coupon was completely covered in a final printing step. To facilitate resistance measurement, the exposed copper wires of the coupon were soldered onto an insulated wire that extended outside the printer,Molex leadpastelasera)b)copper wireFigure 3. a) Solder paste dispensing, and b) laser soldering.1140resulting in an increase of approximately 63 mΩ from the initial temperature reading. The additional wire leads were used to minimize fatigue in the coupon’s wire leads and ensure that resistance values remained within the 2-200 mΩ range to maintain consistent accuracy. Prior to encapsulation, a single soldered wire was randomly selected and measured during the encapsulation process. Resistance measurements were recorded for 10 joints. The entire encapsulation process was recorded on camera to identify any potential causes of failure during the printing process.

Resistance measurements were recorded for 10 joints. The entire encapsulation process was recorded on camera to identify any potential causes of failure during the printing process. After the complete fabrication process, the coupon was allowed to cool down to room temperature over a period of 24 hours. Resistance measurements were then taken to compare against the measurements obtained prior to printing the encapsulation layers. The resistance reading were captured two minutes after clamping the probes in place.top layersepoxy cavityconnectorwire encapsulation layersbase polycarbonateFigure 4. Cross-sectional view of coupon highlighting printing sequences: 1) base polycarbonate, 2) wire encapsulation layers, and 3) top layers.After the completion of the fabrication process, the coupon underwent a 24-hour cooling period to reach room temperature.

Resistance measurements were then conducted to compare them with the measurements obtained prior to printing the encapsulation layers. The measurements were taken two minutes after the probes were clamped.A separate coupon, identical to the one being tested, was also subjected to temperature and resistance measurements before being hand soldered. For these measurements, a DAQ module 9214 from National Instruments in Austin, TX, USA, was employed along with an E-type thermocouple. The thermocouple was encapsulated in epoxy material directly above the solder joint. Temperature and resistance measurements were captured during the “full encapsulation” phase, while the printer was printing over the epoxy cavity.

Temperature and resistance measurements were captured during the “full encapsulation” phase, while the printer was printing over the epoxy cavity. The objective was to analyze the thermal profile and assess the impact of temperature on resistance.Results and DiscussionOut of the 60 connector leads that underwent laser soldering, 48 were successfully connected to copper wire and showed electrical continuity after a single exposure to laser energy. An additional 10 joints required a second exposure to the laser to achieve continuity. In this setup, the desired range for a “good” joint was between 58-65 mΩ. Among the joints created on the first attempt, 45 of them fell within this range, while 8 joints went outside the range due to bridging. However, 6 of those bridge joints were subsequently fixed. After the encapsulation process, a total of 41 joints fell outside the desired range.

However, 6 of those bridge joints were subsequently fixed. After the encapsulation process, a total of 41 joints fell outside the desired range. Following the encapsulation of the 58 joints, 7 wire leads broke, and 4 other leads showed no continuity, resulting in 47 remaining leads for comparison to the pre-encapsulation state. Out of these 47 leads, 34 leads experienced an increase in resistance by 3.83%, with a standard deviation of 4.13%. On the other hand, 13 leads experienced a decrease in resistance by -3.77%, with a standard deviation of 5.49%.1141A BCDE1)2)Figure 5. Resistance across a joint of a coupon undergoing laser soldering. 1) full duration where region A represents the laser heating stage from time 0 to 15 seconds, region B represents the laser cooling stage from time 15 to 30 seconds, and region C is environmental cooling beyond 30 seconds.

2) 30-second timeframe where laser is active: region D corresponds to the laser heating stage (t = 0 – 15 s), and region E is laser cooling stage (t = 15 – 30 s). The resistance highlighted in green was measured during the formation of liquid solder, which attached the lead and the wire characterized by drops in resistance. Also, as the solder transitions from a liquid to solid state (t ≈ 22-25 s), the resistance gradually decreased with power.The joints that successfully connected exhibited an average resistance of 59.93 mΩ with astandard deviation of 6.15 mΩ. Prior to the encapsulation process, 6 joints that had experienced bridging were manually fixed. After encapsulation, the average resistance of the total 47 leads was 61.6 mΩ, with a standard deviation of 4.01 mΩ.

After encapsulation, the average resistance of the total 47 leads was 61.6 mΩ, with a standard deviation of 4.01 mΩ. The two joints that initially showed no continuity maintained their state after encapsulation, and the two joints that were bridged prior encapsulation remained bridged, as evidenced by their low resistance of less than 55 mΩ.The following provides a description of characteristics curves illustrating temperature and resistance. These findings are expected to be particularly valuable due to the consideration of using this hybrid AM approach for on-orbit fabrication. The upcoming descriptions and the data represented in Figures 6-7 can serve as a basis for comparison if experiments are conducted under on-orbit conditions.

The upcoming descriptions and the data represented in Figures 6-7 can serve as a basis for comparison if experiments are conducted under on-orbit conditions. Regarding successful joints achieved through the laser soldering process, the resistance showed the following trends (Figure 5): 1) During the initial 0-7 seconds of increasing heat, leads without initial contact showed continuity or a gradual decrease in resistance as liquid solder formed gradually. 2) Between 7-12 seconds, as the solder reflowed and the laser power (and joint temperature) continued to rise, an increase in resistance was observed. 3) From 12-15 seconds, there was another decline in resistance, indicating the occurrence of wetting between the wire and the lead. 4) Throughout the cooling phase, from 15-30 seconds, the joint’s resistance varied as the laser power decreased. 5) Between 30-120 seconds, the resistance gradually decreased until it stabilized at room temperature.1142abcdefFigure 6.

5) Between 30-120 seconds, the resistance gradually decreased until it stabilized at room temperature.1142abcdefFigure 6. Resistance across a representative joint of a coupon undergoing encapsulation.

Region a was the initial heating phase inside the chamber (t = 0 to 170 s), region b occurred when the build plate lifted and started the first part of encapsulation (t = 170 to 240 s), region c was when build plate lowered, the chamber was open, and epoxy was applied on the cavity and left for 10 minutes (t = 240 to 780 s), region d was when the build plate lifted and waited for the nozzle to start printing (t = 780 to 950 s), region e involved the printing process, where the entire coupon, including the epoxy cavity, was encapsulated (t = 950 to 1020 s), and region e represents the cooling phase of the coupon after the encapsulation process (t =1020 s onwards).After the completion of the coupons’ manufacturing and cooling, resistance measurements revealed two observations on the 9 out of 60 joints that experienced failure: ether continuity was absent, or the resistance increased by approximately 10% compared to the stable nominal value.

However, none of the resistance measurements taken during the encapsulation process for the 10 coupons indicated such issues. This was primarily due to the uneven heating of the printer and the printing location on the build platform which impacted the wire’s heating and prevented the normalization of collected data. Instead, the trends of the measurements were observed as follows (see Figure 6). 1) Initially, when the component was prepared for printing (region a), the lifting of the printing bed caused changes in resistance as the wire configuration shifted and focused heating was applied to the coupon. 2) As the coupon was being printed on (region b), the resistance initially increased until the nozzle printed directly on top of the wire, resulting in a significant rise in resistance.

2) As the coupon was being printed on (region b), the resistance initially increased until the nozzle printed directly on top of the wire, resulting in a significant rise in resistance. It then decreased (highlighted by a red circle), indicating that the wire was heated by the radiation from the nozzle and the deposited material (T = 345 °C), followed by cooling to the chamber environment (145°C). 3) Opening the chamber door (region c) and dispensing epoxy over the joint led to a decrease in resistance as the fluid was at room temperature, cooling the joint and wire. 4) While the curing process was underway, the resistance increased as the epoxy underwent heating inside the chamber. 5) After a period of 10 minutes had elapsed, the printing bed was lifted, and the wire configuration was different (region d).

5) After a period of 10 minutes had elapsed, the printing bed was lifted, and the wire configuration was different (region d). 6) Significant resistance increases were observed when the nozzle passed through the epoxy (region e).1143reflow temperature217⁰Cmax operating temperature85°CabcdFigure 7. Resistance and temperature profile of a hand-soldered coupon undergoing encapsulation for the whole coupon and cavity. In region a, epoxy was dispensed onto the cavity, followed by curing of approximately 10 minutes (t = 0 to 640 s). In region b, the build plate lifted and prepared for the printing phase (t = 640 to 780s). Region c is when printing started from 780 s to 900 s. Region d is after printing, where temperature decreased with resistance (past 900 s).During the process of fabricating the epoxy cavity, the average maximum resistance of 10 joints was recorded as 152.6 mΩ, with a standard deviation of 6.66 mΩ.

These resistance values are higher compared to previous tests due to the increased length of the wires. It is worth noting the remarkable consistency in resistance measurements among the 10 joints at that particular stage, as indicated by a low coefficient of variation (CV = 0.043). Similar consistency was also observed at the completion of printing (end of region e) with an average resistance (µ) of 157.4 mΩ, a standard deviation (σ) of 15.69 mΩ, and a coefficient of variation of 0.099.Regarding the resistance and temperature measurements (see Figure 7), the thermocouple recorded a maximum temperature of approximately 170°C, while the maximum resistance reached around 154 mΩ. At this temperature, the solder joint remained in its solid state. However, any potential failures could have been caused by factors such as vibration, temperature variations, or prolonged exposure to temperatures exceeding the connector’s maximum operating temperature of 85°C.

During this stage, there was no contact between the nozzle and the wires; only vibration occurred due to the rapid movements of the nozzle and the build platform.1144ConclusionThe presented research imparts on the thermal history and resistance behavior of laser soldered joints during their encapsulation using 3D printing. These findings serve as a foundation for future experiments, particularly those conducted under on-orbit conditions. Additionally, this study successfully demonstrated an automated laser soldering solution for fine pitched leads through the implementation of a hybrid additive manufacturing approach for 3D printed electronics. As AM continues to expand its applications in producing innovative devices, this work highlights the integration of multiple technologies, resulting in the creation of coupon-level components that possess physical, mechanical, and electronic functionalities.

The ability to achieve multiple functionalities is crucial for various applications that demand space-saving characteristics, such as satellites.Although the in-situ monitoring method cannot be employed for every solder joint due to its intrusive nature, the methodology presented in this study can be utilized to develop and validate design and process parameters. This includes parameters like epoxy cavity dimensions to minimize undesirable thermal effects, as well as laser power and pulse duration. Additionally, the method proves valuable in identifying the root causes of failure and detecting faulty joints during laser soldering. This enables the opportunity to perform a second laser exposure to repair the joint an enhance overall yield.

This enables the opportunity to perform a second laser exposure to repair the joint an enhance overall yield. During the laser soldering process, the resistance profile revealed the following key observations: 1) an increase in the solder’s molten volume led to a decrease in resistance, 2) the resistance response indicated successful joint attachment, even if visual evidence was challenging to obtain, 3) substantially high resistance response indicated an unsuccessful joint with no attachment. Regarding the temperature profile during encapsulation, it was observed that the temperature at the joint was lower that the solder reflow temperature. Furthermore, resistance measurements taken during the printing process at both stages exhibited increased resistance as the nozzle passed over the wires across each layer.AcknowledgmentsThe research described here was performed at The University of Texas at El Paso (UTEP) within the W.M. Keck Center for 3D Innovation (Keck Center).

Keck Center for 3D Innovation (Keck Center). The authors are grateful to Andrew J. Hall and Jacob I. Rome from Aerospace Corporation for their guidance and support during this research. Also, the authors thank UTEP students: David A. Sepulveda-Navarrete for assisting with machining parameters, Patrick S. Gutierrez for assisting with ultrasonic wire embedding parameters, and Sebastian Oropeza and Anthony Roldan.

SPIE 10698, Space Telescopes and Instrumentation 2018: Optical, Infrared, and Millimeter Wave, 1069826 (6 July 2018)Melanie L., Grande, Daniel A., DeLaurentis, “Evaluating Strategies for an On-Orbit Servicing, Assembly, and Manufacturing Platform with System-of-Systems Methodologies,” AIAA SciTech Forum (January 2021)Michael, Greenstein, “Optical Absorption Aspects of Laser Soldering for High Density Interconnects”, Applied Optics (November 1968): Vol.28, No.21, Pages 4595-4603Ning, Duan, “Comparison of In-Situ Measurement Techniques of Solder Joint Reliability Under Thermo-Mechanical Stresses”, Microelectronics Reliability (October 2014): Volume 54, Issues 9- 10, Page 1753-1757Shemelya, C, "Multi-Functional 3D Printed and Embedded Sensors for Satellite Qualification Structures," IEEE SENSORS (November 2015)1146Xin, Hao, Yu, Xiaoju, Liang, Min, Shemelya, Corey, “3-D Printed Parts for a Multilayer Phased Array Antenna System”, IEEE Antennas and Wireless Propagation Letters (November 2018)1147.

The printing laser beam is diffused and illuminates onto the powder bed’s top surface in various frequencies. A long-wavelength thermal camera monitors the surface temperature history. Insight is provided by a one-dimensional thermal model of the process, which shows the frequency dependence of the surface temperature amplitude and phase on the powder layer thickness and thermal properties. An experiment demonstrates the validity of this model and shows its potential for measuring local powder properties in-situ.1. IntroductionThe laser powder-bed fusion (LPBF) process has potential advantages for manufacturing products with higher dimensional accuracy and complex geometry [1]. In order to improve the quality of the printed products, in-situ inspection techniques have been widely used to study the effects of process parameters such as laser power and scan speed [2,3].

In order to improve the quality of the printed products, in-situ inspection techniques have been widely used to study the effects of process parameters such as laser power and scan speed [2,3]. Although having attracted less attention, powder layer properties also affect the printing process and part quality significantly. For example, powder layer thickness influences the porosity [4,5], microhardness [6,7], surface roughness [8], strength, and elasticity [9,10] of the final part. The powder bed also acts as a support for overhang and other complex structures. The surface qualities of these structures are limited by the thermal properties of the powder bed.

The powder bed also acts as a support for overhang and other complex structures. The surface qualities of these structures are limited by the thermal properties of the powder bed. In order to improve the finish and accuracy of the surface, the laser process parameters need to be optimized based on the knowledge of the local powder’s thermal conductivity [11].The powder layer thickness can be monitored by using high-resolution optical imaging [12], optical coherence tomography (OCT) [13] or fringe projection [14]. The laser displacement approach has also been used to measure the local powder thickness [15]. However, these methods can only capture information about the powder bed surface profile. The thermal conductivity of metal powders has been studied by using ex-situ hot wire [11], ex-situ laser flash [16], and in-situ reﬂectance measurement [17]. These techniques cannot obtain the powder geometry in-situ.

These techniques cannot obtain the powder geometry in-situ. It is much more convenient to measure the powder’s layer thickness and thermal properties at the same time.Infrared thermography (IRT) is one of the nondestructive testing (NDT) techniques using the diffusion of heat through a material to evaluate the subsurface cracks or defects and coating1402thicknesses [18]. Compared with other NDTs like hammer sounding and microwave radar, IRT has drawn more attention due to its advantages of real-time, contactless and wide-area measurements [19]. By using IRT, both thickness [20] and thermal properties [21] can be measured. There are two different types of IRT including passive and active, which are based on different types of heating sources [22].

There are two different types of IRT including passive and active, which are based on different types of heating sources [22]. The passive IRT was used to measure the powder layer’s thickness in our previous work [23], which used the printed part beneath the powder layer as the heating source.This paper investigates an active IRT based technique to estimate the powder layer’s thickness and thermal properties. The printing laser serves as the heating source. By controlling the printing laser exposure time and hatch pattern, the powder layer is heated by absorbing the laser’s periodic illumination with respect to time. The temperature history of the powder’s top surface is observed with a long wavelength infrared (LWIR) camera. After analysis of the temperature history, the transient thermal response is correlated with the powder thickness and thermal properties. A validating experiment is conducted for this correlation.

After analysis of the temperature history, the transient thermal response is correlated with the powder thickness and thermal properties. A validating experiment is conducted for this correlation. A heat transfer model (1D) is established to compare analytical predictions with experimental results.2. Experimental SetupFigure 1 illustrates the experimental setup with a long-wavelength infrared (LWIR) camera (FLIR A655sc) integrated with a commercial LPBF machine (Renishaw AM250). A ZnSe window on the chamber allows the LWIR camera to observe the build plate from 15° off-normal. The LWIR camera has the resolution of 640×480 pixels, corresponding to a spatial resolution of 325 µm for this configuration. The spectral range of this LWIR camera is 7.5-14.0 µm. The noise- equivalent temperature difference (NETD) is 30 mK and the frame rate is 200 Hz (with zoom in 640×120 pixels).

The spectral range of this LWIR camera is 7.5-14.0 µm. The noise- equivalent temperature difference (NETD) is 30 mK and the frame rate is 200 Hz (with zoom in 640×120 pixels). The camera reports the temperature of the scene assuming an emissivity of 0.95 (gray-body approximation), which is not corrected for transmission through the window. After transferring the temperature from time domain to frequency domain, the amplitude can be separated from the phase information. This removes requirements for emissivity information and non-uniformity calibrations, significantly simplifying the instrumentation. The laser is expanded by an optical diffuser (DG100×100-220, Thorlabs) and fixed 30 cm above the surface of the powder bed.Laser and Beam Delivery OpticsScannerLWIR CameraWindowBuild ChamberLaserZnSeOptical diffuser15°CFront windowPowderPrinted BlockδyCatchmentBackBuild PlatezElevatorFigure. 1.

1. Illustration of the LPBF experimental setup including a LWIR camera.1403In order to collect the temperature histories for various known powder thicknesses, simple gage specimens are printed with the Renishaw AM250. They consist of 7 cuboid blocks (5×5 mm2) with different elevations, h, from a printed datum plane. After the specimen is printed, the wiper (recoater) then spreads new powder over the build plate including the specimen. This gives a powder layer thickness that varies spatially with the height of the individual blocks. The powder layer thickness is δ=h0-h, where h0 is the height above the datum of the build plane. In this experiment h0=1 cm so that the highest specimen has δ=100 µm of powder above it. The heights of the gage specimens and corresponding powder thicknesses are listed in Table 1.

In this experiment h0=1 cm so that the highest specimen has δ=100 µm of powder above it. The heights of the gage specimens and corresponding powder thicknesses are listed in Table 1. The temperature of the build plate is constant at 50 °C.Table 1 The heights of seven different blocks and corresponding powder thicknessesn h [mm] 1 2 3 4 5 6 79.90 9.85 9.80 9.75 9.70 9.65 9.60δ [µm] 100 150 200 250 300 350 400Figure 2 shows the temperature history of a single pixel for the 200 W laser of AM250 illuminating the sample 100 times at the frequency of 0.25 Hz, using 304L stainless steel powder with different layer thicknesses. The powder used has particle diameters ranging between 15 and 40 µm, with a mean diameter 25 µm. The laser illuminating has a square wave. As shown in Fig. 2, from t ≈ 8 s the temperature of the powder starts to increase from around 50 °C and oscillates at 0.25 Hz with the laser’s periodic illumination.

As shown in Fig. 2, from t ≈ 8 s the temperature of the powder starts to increase from around 50 °C and oscillates at 0.25 Hz with the laser’s periodic illumination. The laser is turned off after 100 cycles at t ≈ 408 s, and then the powder temperature begins to cool down.40002004006005052545658Temperature, T [°C]Time, t [s]Powder thickness, δ [μm]100200(a)10038638839039252545658Temperature, T [°C]Time, t [s]Powder thickness, δ [μm]200400(b)Figure. 2. Temperature history of a single pixel for (a) a periodic heating process at 0.25 Hz for 100 times and the dash marked portion is shown in (b).The periodic temperature history can be fitted with a Fourier series by using linear leastsquare regression as follows:1404()()()()11122221===+++=++−=−NDnnnnNDnnnnnnTCtTcosπftφCtαcosπftβcosπftfnfwhere T is temperature, t is time, Tn and φn are the amplitude and phase of the temperature at frequency fn.

Since the laser heating source has a square wave, the frequency fn is an odd integer multiplying the fundamental frequency f1 (e.g., 0.25 Hz). tD is the power function of t with index D. C, D, αn, and βn are fitting coefficients. Then the amplitude and phase of the temperature are221−=+=nnnnnnTαββφtanαAn example fitting result is shown in Fig. 3. After fitting the temperature history by using Eq. 1, the values of C, D, αn, and βn are determined. Then the phases for different powder thicknesses at various frequencies can be obtained.Fitting by N=5300350400525354Temperature, T [°C]Time, t [s]ExperimentFigure. 3. Comparison between experimental data with the fitting result for δ = 100 µm.Figure 4 shows the phases for different powder thicknesses at the first (fundamental) frequency of the square wave, f1 = 0.25 Hz.

It is obvious that the phase can be correlated to powder thickness with a monotone relationship: the larger the powder thickness, the smaller the phase angle. This relationship can be used to predict the powder thickness using the phase difference between the periodic laser heating and the measured surface temperature.100200300400-2.4-2.2-2.0-1.8Phase, φ [rad]Powder thickness, δ [μm]Frequency, f1= 0.25 HzFigure. 4. Experimental correlation between phase and powder thickness with the fundamental frequency f1 = 0.25 Hz.1405(1)(2)3. 1D Heat Transfer Model and Finite Difference SolutionThe powder surface temperature history can be predicted by modeling the thermal diffusion through the powder from the periodic laser heating. Since the thickness of powder layer is much smaller than the lateral dimensions of the blocks (5 mm), it is reasonable to model the problem as 1D heat transfer in z-direction.

Since the thickness of powder layer is much smaller than the lateral dimensions of the blocks (5 mm), it is reasonable to model the problem as 1D heat transfer in z-direction. The powder is evaluated as a continuous medium, with its thermal properties being temperature independent. A schematic diagram of this model is shown in Fig. 5(a), which shows a powder layer of thickness δ is in contact with a printed block. A diffused laser with illumination heat flux qlaser in a square wave is emitted on the top surface of the powder, and it is assumed that the heat from the laser is completely absorbed by the powder (100% absorption). A convection boundary condition is applied with h = 15 W/m2K. The printed block is in perfect contact with the build plate, whose temperature is constant at 50 °C.

A convection boundary condition is applied with h = 15 W/m2K. The printed block is in perfect contact with the build plate, whose temperature is constant at 50 °C. The environment temperature is 30 °C.The heat transfer problem as described is governed by the following equation:22iiikTTtρcz=with the boundary conditions()()00,30,50===−−===plaserzTkhTztTqzTCTzLtCand the initial condition()050==Tz,tCThe thermal properties in Eq. (3) depend on the form of material, which may be either powder or printed block (denoted by the subscript, i, which is p for powder and b for block). k, ρ and c are the effective thermal conductivity, density and specific heat of the material, respectively.

5. The 1D heat diffusion model: (a) model definition, (b) laser heating intensity and (c) powder surface temperature response.Since a square wave can be represented by a Fourier series with infinite terms in frequency domain, the linear heat transfer equation (3) also can be solved in frequency domain. The finite difference solution of the 1D heat transfer equation in frequency domain is shown in Fig. 6 for varying powder thickness. Figure 6(a) shows the phase vs. frequency relationship for different powder thicknesses from 100 µm to 400 µm. The dispersion phases for f1 = 0.25 Hz are shown in Fig. 6(b), which also has a monotone relationship similar with the experimental result in Fig.

The dispersion phases for f1 = 0.25 Hz are shown in Fig. 6(b), which also has a monotone relationship similar with the experimental result in Fig. 4.2002500.250.500.751.001.251.501.752.00-2.4-2.2-2.0-1.8100350150400Phase, φ [rad]Frequency, f [Hz]Powder thick δp [μm]Powder thermal conductivitykp 0.269 W/mKPowder densityρp 4,818 kg/m3300(a)100200300400-2.2-2.0-1.8Phase, φ [rad]Powder thickness, δ [μm]Powder densityρ 4,818 kg/m3Thermal conductivityk 0.269 W/mKf=0.25 Hz(b)Figure. 6. The finite difference solution of the 1D heat diffusion model in frequency domain with varying powder thickness (a) in frequency domain and (b) for f1 = 0.25 Hz.The effect of powder thermal conductivity on the temperature phase is also studied, and the result is shown in Fig. 7(a). This result denotes that all the phase values are dispersive in the frequency range of 0.25-2.00 Hz. At the frequency f1 = 0.25 Hz, as shown in Fig.

7(a). This result denotes that all the phase values are dispersive in the frequency range of 0.25-2.00 Hz. At the frequency f1 = 0.25 Hz, as shown in Fig. 7(b), the monotone relationship between the phase and the powder’s thermal conductivity shows the potential for using the phase to monitor in-situ the change in local powder thermal conductivity during the LPBF process.14070.269140.250.500.751.001.251.501.752.00-2.4-2.2-2.0-1.823 Phase, φ [rad]Frequency, f [Hz]Powder thick, δ 100 μmPowder density, ρ 4,818 kg/m3Thermal conductivity, k [W/mK]01234-2.4-2.2-2.0-1.8Phase, φ [rad]Thermal conductivity, k [W/mK]Powder thick, δ 100 μmPowder density, ρ 4,818 kg/m3f = 0.25 HzFigure. 7.

7. The finite difference solution of the 1D heat diffusion model in frequency domain with varying powder thermal conductivity (a) in frequency domain and (b) in 0.25 Hz.It should be noted that the square wave of laser heating results in the surface temperature that can be converted to a Fourier series with infinite terms, with each term corresponding to a different phase. However, the amplitudes of higher-order terms are getting smaller and smaller. Because of their small amplitudes, plus the fact that they may be overwhelmed by the camera noise and powder temperature fluctuations in a real environment, these terms can be ignored in practical implementation.4. ConclusionThis paper demonstrates the potential of using in-situ active thermography to predict the powder layer thickness and to estimate local powder layer thermal properties during the LPBF process.

We have achieved this goal by using an optical diffuser to expand the size of laser beam in a commercial LPBF machine (Renishaw AM 250). A LWIR camera is used to measure the surface temperature resulted from square-wave laser illumination. By analyzing the powder surface temperature history, a monotonic relationship between powder thickness and the phase between surface temperature and laser illumination is obtained at a low fundamental frequency (f1 = 0.25 Hz). This phase relationship can be used to predict the powder thickness during the LPBF process. The experimentally obtained relationship agrees with the relationship obtained from simulations by solving the 1D heat transfer model numerically in frequency domain.

The experimentally obtained relationship agrees with the relationship obtained from simulations by solving the 1D heat transfer model numerically in frequency domain. Besides the prediction of powder thickness, the solution of the 1D heat transfer model also demonstrates the potential of using the active thermography with a LWIR camera to monitor change in powder thermal conductivity during the LPBF process.Our future work will include integrating the active thermography method with a machine learning method for improved predictions. Machine learning could potentially result in better correlations of the phase obtained from the powder surface temperature measurement with powder thickness and also the powder’s thermal properties. This is because the large mount of data associated with the large number of Fourier series terms can be incorporated into machine learning without any difficulty.1408Reference1.

It enables the collection of relevant measurement data at previously inaccessible component regions, which are needed for the determination of maintenance strategies. Therefore, the following research deals with the development of a concept, with which weldable strain gauges (WSG) can be weld-on by the laser-scanner-system of a PBF-LB/M system to the manufactured component. The examinations show that the arrangement of so-called weld seams has a decisive influence on a correct metal carrier connection of the WSG. It can be deduced from the study that WSGs can be integrated within components manufactured by PBF-LB/M and completely welded to it with the laser-scanning-system.

It can be deduced from the study that WSGs can be integrated within components manufactured by PBF-LB/M and completely welded to it with the laser-scanning-system. Following measurement validations demonstrate that the WSG provides complete and correct measurement data and can therefore be considered as successfully connected to the component.Keywords: Laser-Based Powder Bed Fusion, Metal Additive Manufacturing, Strain Gauges, Smart Parts, InconelIntroductionAdditive Manufacturing (AM) in general describes technologies to produce components by successive addition of material and are usable for a wide range of applications in industry. Present statistics indicate a vigorously growing of AM industry over the next several years (Compound Annual Growth Rate of 23.3 %), because of its potential to generate a variety of economic benefits, like manufacturing of individual products and efficiency improvements in production [1,2].

The expansion of the current state of the art is therefore of great technical and economic interest.Recent studies investigating developments and white spots in additive manufacturing identified the embedding of sensors in AM components as one of the most important trends with a high need for research [3–5]. The benefit of the layer-by-layer process is the possibility of sensor integration at any time during the process and at any area within the manufactured part [6–9]. Through the surrounding component, an integrated sensor can be protected from external environmental influences, is not visible from the outside and does not influence its geometrical shape. In addition1878structural weakness of the part through sensor integration can be minimized and therefore a high functional density be realized. These advantages allow the acquisition of relevant measurement data at previously inaccessible component sections.

These advantages allow the acquisition of relevant measurement data at previously inaccessible component sections. This approach can potentially improve the implementation of predictive maintenance methods and capture more accurate process data that enables advanced process control for processing machines in the chemical industry, for example.The integration of sensors during the metalworking laser-based powder bed fusion (PBF-LB/M) [10] process requires an extension of the existing manufacturing process. This involves a process interruption at any cross section layer of the component to be manufactured, while at the same time it must correspond to the top of a cavity manufactured within the component. In addition, the loose metal powder inside must be removed, the sensor integrated, and the manufacturing process continued. Sensor integration during the PBF-LB/M process involves numerous challenges, which must first be considered and overcome.

Sensor integration during the PBF-LB/M process involves numerous challenges, which must first be considered and overcome. On the one hand, the sensor to be integrated must be able to withstand the high temperature gradients during the manufacturing process and must not be damaged by them. On the other hand, a sufficient connection to the carrier component must be ensured so that the sensor supplies correct and usable measurement data. Care must also be taken to ensure that the internal cavity in which the sensor is integrated does not have any negative effects on neither the component strength nor the accuracy of the sensor.The strain gauges are considered promising for numerous measurement tasks and are thus among the most widely used sensors for component monitoring. Strain gauges are sensors for measuring stretching and compressing deformations. They are also suitable for stress analysis.

Strain gauges are sensors for measuring stretching and compressing deformations. They are also suitable for stress analysis. For this, the mechanical stresses prevailing in an object are determined from the measured deformations by applying the corresponding material laws. Depending on the strain gage design and the resulting measurement requirements, the application can be carried out in various ways. In addition to conventional adhesive bonding, which is prevalent in the PBF-LB/M process due to the metal powder particles present and which can severely impair this process, it is also possible to weld the strain gauges to the component surface using an appropriate metal carrier and thus achieve a material bond. Numerous factors must be taken into account here, which are first determined and then qualified.

Numerous factors must be taken into account here, which are first determined and then qualified. Accordingly, weldable strain gauges are to be integrated in components during PBF- LB/M in components consisting of the nickel-chromium alloy by means of laser welding with the existing laser-scanner system.Within this study the possibility of integrating weld-on strain gages during PBF-LB/M is to be explored. In addition to the strain gage and the cavity design, particular attention will be paid to the weld seam design and the actual welding process. For this, the necessary parameters for welding the strain gauge are first defined and then qualified in several tests. This is to ensure that the strain gage has a strength bond to the substrate material. In the following, the determined process parameters for the strain gauge connection are also applied during PBF-LB/M and a strain gauge is integrated into a carrier component.

In the following, the determined process parameters for the strain gauge connection are also applied during PBF-LB/M and a strain gauge is integrated into a carrier component. Subsequently, a vibration test method is carried out in which the forced frequencies of the integrated strain gauge are recorded and converted into measurement data in order to investigate the bonding connection of the strain gauge. In the outlook, the possibilities of automated sensor integration are described.1879Materials and Methods Powder Material for PBF-LB/MIn this research, nickel-chromium powder (NickelAlloy IN718) provided by EOS GmbH [11] was processed by an AconityONE system with variable laser power and scan speed. For the manufacturing of the components the process parameters were adapted for sufficient material density.

For the manufacturing of the components the process parameters were adapted for sufficient material density. Therefore the layer thickness (30 µm), a nominal focused spot size (80µm), hatch distance (100 μm) and rotation angle 67° of scan strategy (meander) is set to a constant value. Thus results in material density of about 99.85 ± 0.05 % for a laser power of 300 W and a scan speed of 700 mm/s of the manufactured components were achieved. Process gas environment is formed by argon at a constant build temperature (no base plate heating). The powder was recycled through sieving during the several build jobs to ensure uniformity using a sieve with 68 μm mesh size.Weld Seam ShapeWith regard to the sensor connection by means of laser welding with the laser-scanner system, special care must be taken to ensure that the sensor component is fully connected and thus sufficient measurement accuracy is guaranteed.

Consequently, an inadequate strain gauge connection can result in corresponding measurement deviations. For the used preparation software (Autodesk Netfabb) there is no possibility to create single weld lines during the PBF-LB/M process. Therefore, the required weld seams are realized by a contour exposure as shown in Fig. 1. This results in a weld seam length as well as a weld seam width which can be changed subsequently and thus define the weld seam dimensions.Fig. 1 Schematic representation of a weld seamWeld Seam Analysis (Feasibility Study)The aim of this research is to investigate the effect of process parameters such as laser power and scan speed on the melt-pool expansion of solid components during the PBF-LB/M process.

Therefore, the process parameters can be used for the subsequent investigation of the metal carrier welding and ensure a firmly bonding and low thermal impact on the sensor component.Process ParametersFor this procedure, the process parameters laser power, scan rate and focus shift were considered because these three have a decisive influence on the expansion of the melt-pool during PBF-LB/M [12]. In preliminary investigations the influence of laser power and scan speed on the material density and therefore the melt-pool expansion during PBF-LB/M were considered. This means that the largest melt pool expansion can be expected with increasing laser power at reduced scan speed.1880Focus shift can be described as a distance between the build plate with a coated layer of powder and the focus plane of the laser beam [13]. The system is typically calibrated so the laser focal plane coincides with the zero-position (Fig. 2).

The system is typically calibrated so the laser focal plane coincides with the zero-position (Fig. 2). In a PBF-LB/M system the laser focus shift measures how far the build plate is moved above (negative shift) or below (positive shift) the zero- position of the build chamber. Defocusing the laser via focus shift occurs with no changes to the laser optics, so it results in the same linear energy density but a larger spot size, and therefore has the effect of changing the area and volumetric power density.Focus radius in µm 150 50 0 50 1500.2m m n i0t f i h s0.2s u c o F0.40.6Beam causticFig. 2 Principle representation of the focus shift at zero-positionTest SpecimenThe test specimen shown in Fig. 3 were manufactured using an AconityONE system with the defined process parameters for a material density of 99.85 % (±0.05 %).

3 were manufactured using an AconityONE system with the defined process parameters for a material density of 99.85 % (±0.05 %). Then the loose metal powder around the manufactured specimen is removed from the surfaces within the PBF-LB/M machine.A-A (Longitudinal)Top viewDetail view) s i x a - z ( n o i t c e r i d d l i u Bm m 2.00.40 mmm m 0 0.5 1) s i x a - z ( n o i t c e r i d d l i u Bm m 0 0.0 10.18 mm) s i x a - z ( n o i t c e r i d d l i u Bk c a r T dwTrackTest specimenWeld trackWeld seamFig. 3 Test specimen for the weld seam analysis with depth of the welding track dTrack and width wTrack1881Here it is ensured that the test specimen surfaces do not contain any metal powder that can influence the expansion of the melt-pool.

Subsequently, the PBF-LB/M system is again flooded with the process gas (argon) and the focus of the laser scanner system is focused on the lowest test specimen surface (lowest step), whereby the weld seams are applied with the corresponding laser powers and scanning speeds. The "steps" of the test specimens shown in Fig. 3 allow that the focus shift of the laser scanner system does not have to be changed for the weld seam application. Then the substrate plate can be removed and the weld seams applied to the test specimens can be examined.Metal carrier welding Process parametersFor the metal carrier welding the process parameters of the previous investigation were used. Therefore the laser power and scan speed were used to regulate the melt-pool expansion and thereby the thermal impact on the metal carrier. However further process parameters were defined to ensure a firmly bonding of the metal carrier on the component surface.

However further process parameters were defined to ensure a firmly bonding of the metal carrier on the component surface. In previous research it could be found out that the weld seam arrangement has a significant influence on the metal carrier connection using a laser-scanner system [14].Test SpecimenFor the metal carrier connection test specimens were designed, which have a cavity construction in which the metal carrier can be inserted. The exact dimensions are shown in Fig. 3. The test specimens were manufactured using an AconityONE and SLM 125 system with the defined process parameters for solid material components and then the surrounding metal powder was removed from the test specimens within the PBF-LB/M system.

Afterwards the metal carriers are manually deposited on the cleaned cavity surfaces and the laser beam melting system was flooded again with the process gas, whereby the welds with the defined process parameters were applied in the transition area of the cavity surface and the metal carrier. The substrate plate was then removed and the weld seams applied to the test specimens were examined.A-A (Longitudinal)Top viewDetail view) s i x a - z ( n o i t c e r i d d l i u B) s i x a - z ( n o i t c e r i d d l i u BWeld seam) s i x a - z ( n o i t c e r i d d l i u Bk c a r T dwTrackTest specimenWeld trackMetal carrierFig. 4 Test specimen for the metal carrier welding1882Metallographic methodsIn this research, the specimens were prepared using standard metallographic methods. For this purpose, the specimens are cut to prepare sections as shown in Fig. 3 and Fig. 4 (see A-A and Detail view). Then the resulting relevant specimen segments were embedded.

For this purpose, the specimens are cut to prepare sections as shown in Fig. 3 and Fig. 4 (see A-A and Detail view). Then the resulting relevant specimen segments were embedded. Afterwards the specimens were grounded and polished. For contrast generation, the resulting specimen sections must be etched. For this purpose, the etchant "Adler-Etching" (Etching time = 50 seconds) was used, which is mainly used for weld seam inspection of nickel-chromium alloys. Subsequently, light microscopy was performed with different magnifications, which allowed the weld seams and melt- pool expansions to be analyzed.Results and Discussion Weld seam analysisFor the weld seam analysis in addition to the laser power and scan speed the focus shift was also considered. The process parameter combination of laser power and scan speed can be described by the linear energy. Therefore both the laser power and scan speed were adjusted so that defined linear energies were obtained, cf.

Therefore both the laser power and scan speed were adjusted so that defined linear energies were obtained, cf. Table 1.Table 1 Process parameters for weld seam analysisTest combinationsLaser power in WScan speed in mm/sLinear energy in J/mm0.50 0.75 1.00 1.25 1.50 0.50 0.75 1.00 1.25 1.50Focus radius in µm 150 50 0 50 150Focus radius in µm 150 50 0 50 150Focus radius in µm 150 50 0 50 1500.20.20.2000m m n i0.2m m n i0.2m m n i0.2t f i h s0.4t f i h s0.4t f i h s0.4s u c o F0.6s u c o F0.6s u c o F0.60.80.80.81.01.01.0Focus shift: 0.00 mmFocus shift: -0.20 mmFocus shift: -0.40 mmBeam caustic1883For a scanning speed of 200 mm/s (const.), the individual laser powers were changed in such a way that the individual linear energies 0.50 J/mm ≤ EL ≤ 1.50 J/mm (∆EL = 0.25 J/mm) resulted.

In addition, for a laser power of 300 W (const.), the corresponding scan speeds were also changed so that the same linear energies resulted. Thus it can be investigated to what extent both the laser power and the scanning speed influence the resulting melt pool expansion. For the weld seam inspection, the melt pool expansion of the same linear energies can also be examined under different laser powers and scanning speeds. For the weld seam investigation, the focus shift was changed in addition to the listed parameter combinations of laser power and scan speed, which according to Beyer should lead to an increased melt pool expansion [12].Weld seam expansionFig. 5 initially shows that the weld track depth was not significantly influenced within a defined laser power and scanning speed. Slightly different weld track depths were due to measurement uncertainties caused by contrast differences in the light microscope images.Fig.

Slightly different weld track depths were due to measurement uncertainties caused by contrast differences in the light microscope images.Fig. 5 Weld track depths at different laser powersOefele describes a so-called tolerance field for the laser scanner system, which results from the Rayleigh length of the laser beam [15]. Within this tolerance field, the weld track depth is not influenced by the focus position, since the laser beam profile hardly expands. For an increased Rayleigh length, the tolerance field of the laser beam increases at the same time, which is why the focus position should lie outside this tolerance field in order to influence the weld track depth significantly. No beam measurement documents were available at the time of this study for the1884AconityONE PBF-LB/M system, which is why the Rayleigh length and the resulting tolerance field are not known.

No beam measurement documents were available at the time of this study for the1884AconityONE PBF-LB/M system, which is why the Rayleigh length and the resulting tolerance field are not known. Consequently, the focus position can be considered insignificant for the weld pool expansion.For the laser power PL = 300 W, decreasing scan speeds and the associated increased linear energies also resulted in increased weld track depths, which differ significantly from those of the lower linear energies. This can also be observed for the scan speed vS = 200 mm/s with increasing laser power. Accordingly, both the laser power and the scanning speed can be regarded as decisive process parameters for the melt pool expansion and the resulting weld track depth. With regard to the linear energies, it emerges that the weld track depth also increases with increasing line energy.

With regard to the linear energies, it emerges that the weld track depth also increases with increasing line energy. It should be noted that different weld track depths were produced for identical linear energies under varying laser power and scan speed. Consequently, it can be concluded that there is a non-linear relationship between the energy input and the resulting weld pool expansion.Weld seam integrityFor weld seam inspection, it is recommended to view the light microscope images of the weld seams in addition to the weld track measurement. In addition to the weld shape, this also allows the weld integrity to be assessed. Fig. 5 showed the light micrographs of the weld seams for the linear energies EL = 0.50 J/mm and EL = 0.75 J/mm and a focal position of zF = 0.00 mm. Due to the low metal substrate thickness of the LS31HT strain gauge, weld track depths of 0.15 mm to 0.50 mm are considered sufficient for sensor attachment and were examined in more detail accordingly.

Possible contrast differences are due to a slightly different etching time of the test specimens.With regard to weld integrity, it can be seen that with increasing laser power and simultaneously decreasing scanning speed and resulting capillary formation within the weld track, limited material separations (microcracks) occured. According to Anik and Dorn, hot cracks can be assumed on the basis of their shape and position (Fig. 5), which occur just below the solidus temperature when the molten metal solidifies [16]. Due to the fact that the molten metal pushed out in front of the crystallization front forms isolated micro segregations and thus has a lower solidus temperature than the already solidified grains, the molten metal can no longer sufficiently compensate for the shrinkage that occurs. If the molten metal no longer flows sufficiently and the deformation capacity is exceeded, microcracks will form within the weld track.

If the molten metal no longer flows sufficiently and the deformation capacity is exceeded, microcracks will form within the weld track. According to Anik and Dorn, nickel- chromium alloys have an increased tendency to hot cracking during the welding process. Consequently, the microsegregations of the molten metal should be reduced, which can ensure weld integrity. As a result of the lowest possible temperature gradients during the welding process through suitable process parameters (laser power and scanning speed), the microsegregations and thus the microcracks can be avoided. In addition, the shrinkage hindrance can be kept as low as possible by a suitable welding sequence, which also avoids the microsegregations.Ultimately, it can be determined that different weld track depths result from different laser powers and scanning speeds and can accordingly be regarded as decisive for the melt pool expansion.

Due to the fact that the sensor component needs not be damaged and the strain gauge LS31HT has a low metal substrate thickness, the welds resulting from the laser powers PL = 100 W and PL = 150 W with the corresponding scan speed vS = 200 mm/s (test combination 1.1 and 1.2) are considered sufficient for the subsequent sensor connection. In addition, temperature gradients can thus be kept1885low and weld integrity can be ensured. For the focal position, it emerges that focal positions -0.40 mm ≤ zF ≤ 0.00 mm do not significantly influence the generated weld shape and thus focal position zF = 0.00 mm is specified for the sensor connection.Metal carrier connection Preliminary testFor the preliminary test, the laser powers PL = 100 W and PL = 150 W already determined in the previous weld inspection were used with an associated scan speed vS = 200 mm/s.

The welding track depths thus achieved are considered sufficient to melt both the metal platelet and the underlying metal surface and to form a coherent metal melt. With the relatively low laser powers and the associated lower temperature gradients, the weld integrity can also be ensured. With regard to the weld seam investigations, it emerged that the focal position did not demonstrably influence the melt pool expansion, as a result of which a focal position zF = 0.00 mm was provided for the metal platelet bond. Following the VDI/VDE Guideline 2635-2 [17,18] and the manufacturer's data sheet of the LS31HT strain gauge, various weld seam arrangements were defined for the laser- scanner system. Accordingly, both weld seam arrangement I and weld seam arrangement II were investigated. Table 2 lists the resulting parameter combinations for the metal plate attachment. According to VDI/VDE Guideline 2635, the metal platelets should first be provided with two tack welds (T).

Table 2 lists the resulting parameter combinations for the metal plate attachment. According to VDI/VDE Guideline 2635, the metal platelets should first be provided with two tack welds (T). This is to ensure that the metal platelets can be provisionally tack-welded to the cavity support surface and that the subsequent welds can be applied correctly. Starting from the tack welds, the subsequent welds were applied according to Table 2 with the intended weld numbering on both the metal platelet and the cavity support surface, which can be expected to provide a coherent metal melt. With the Autodesk Netfabb software used, the weld width cannot be reduced arbitrarily, so the smallest possible weld width 0.18 mm was examined first.

With the Autodesk Netfabb software used, the weld width cannot be reduced arbitrarily, so the smallest possible weld width 0.18 mm was examined first. The complete weld seam dimensions are listed in Table 2 (Detail View).Table 2 Process Parameter for Preliminary TestLaser power in WScan speed in mm/sFocus shift in mm Weld seam width in mm0.00 0.002.00 2.00T8910 11121314T24681012140.30 mm0.18 mmT1234567T135791113m m 0 61m m 0 3.2III2.00 mmDetail viewFor the metal platelets, considerable material distortion can initially be demonstrated (see Fig. 6). According to Langhorst, the material distortion can be attributed to increased temperature gradients and an inhomogeneous temperature distribution within the metal platelet [19]. As a result of the finite conduction conditions of the inhomogeneously heated metal platelet, welding residual stresses result, which caused the material distortion.

As a result of the finite conduction conditions of the inhomogeneously heated metal platelet, welding residual stresses result, which caused the material distortion. Especially for metal platelets, the welding- induced material distortion can be caused abruptly by material imperfections within the metal platelet and can hardly be controlled due to the unclear material situation.1886Fig. 6 Weld track depths at different laser powers (width of the metal plate = 10 mm)For the weld seam arrangement I and II, with which the laser-scanner system applied the weld seams, it can thus be determined that the successively applied weld seams cause considerable temperature gradients within the metal platelet and the associated material distortion The tack weld seams, which were intended to adhere the metal platelet to the cavity support surface, were unable to prevent the material distortion.

From the metal plate images, it appears that the test combinations 1.1 and 1.2 (PL = 100 W and vS = 200 mm/s) were not completely welded through and thus did not result in sufficient melt pool expansion to bond the metal platelet to the cavity support surface. Consequently, a laser power PL = 100 W with a corresponding scan speed vS = 200 mm/s can be regarded as insufficient for the metal support bonding. For the test combinations 2.1 and 2.2, it can be seen that laser power PL = 150 W with an associated scan speed vS = 200 mm/s can be considered sufficient to melt both the metal platelet and the cavity support surface and form a cohesive metal melt.

It should be noted here that sufficient metal platelet bonding cannot be demonstrated due to the resulting material distortion.With regard to the VDI/VDE guideline 2635-2 and the manufacturer's data sheet of the LS31HT strain gauge, a metal carrier connection with the spot welding process is recommended, whereby the metal carrier is pressed onto the metal surface with a defined force and thus the material distortion can be compensated. For metal substrate bonding with a laser-scanner system, material distortion can only be prevented by suitable process parameters and the associated, sufficient temperature distribution. Consequently, a weld seam arrangement must be determined by which the metal substrate is heated uniformly and thus the material distortion is prevented.1887Modified testFor metal platelet bonding, the laser powers PL = 100 W and PL = 150 W, which have already been investigated in detail, were used with the associated scan speed vS = 200 mm/s.

This means that the process parameter combinations to be investigated can be kept low. An adapted weld seam arrangement was determined for the metal substrate connection, through which the metal plate is heated as uniformly as possible and material distortion is thus prevented. Previous research has already shown that, in addition to laser power and scanning speed, multiple weld seam exposure can be regarded as an elementary process parameter [14]. Due to the fact that the metal platelets could not be completely welded with a laser power PL = 100 W and the corresponding scan speed vS = 200 mm/s, the metal platelets were accordingly also exposed several times. Consequently, different weld exposures result for the applied weld seam arrangement III, IV and V, which are listed in Table 3. Due to the droplet formation of the applied welds, the weld width was additionally defined as a process parameter for the metal platelet bond.

Due to the droplet formation of the applied welds, the weld width was additionally defined as a process parameter for the metal platelet bond. In addition to the weld seam width 0.18 mm already investigated, the weld seam width 0.36 mm was therefore used. Consequently, the aim was to ensure that the adjacent weld tracks do not influence each other and that the drops of the welds can thus be avoided. Table 3 lists the modified process parameters for the metal platelet attachment.Table 3 Process Parameter for Modified TestLaser power in WScan speed in mm/sFocus shift in mm Weld seam width in mm0.00 0.00 0.00 0.002.00 2.00 2.00 2.00The process parameter combinations show that only individual process parameters were changed, which allows the various process parameter influences to be assessed. With regard to VDI/VDE Guideline 2635-2, the weld seams (1) are first applied by the laser scanner system.

With regard to VDI/VDE Guideline 2635-2, the weld seams (1) are first applied by the laser scanner system. As a result, the metal platelets were already evenly bonded to the cavity surface and material distortion was reduced. Starting from the weld seams (1), the subsequent weld seams were applied according to Table 3 with the intended weld seam numbering both on the metal platelet and on the cavity contact surface, whereby a coherent metal melt is expected. In this case, the weld seam arrangement was specifically defined to generate the largest possible spacing of the same weld seam numbering.The metal plate bonding was performed on the SLM125 laser beam melting system, as mentioned earlier. For the modified test, the same test specimens were manufactured as before with the process1888parameters mentioned for the NickelAlloy IN718 metal powder.

For the modified test, the same test specimens were manufactured as before with the process1888parameters mentioned for the NickelAlloy IN718 metal powder. After the metal plates have been inserted, the PBF-LB/M system was flooded again with the process gas and the weld seams applied according to the weld seam numbering with the process parameters defined in Table 3.For the test specimen images, it can be seen that for a laser power PL = 100 W (Test combination 1.1), there is a discernible material distortion of the metal plate. The metal plate was therefore not completely bonded to the cavity contact surface. The decisive factor here is that, due to the laser power PL = 100 W with the associated scan speed vS = 200 mm/s, insufficient melt pool expansion of the applied weld seams resulted. Consequently, no coherent metal melt formed, which sufficiently bonds the metal plate to the cavity support surface and counteracts material distortion.

Consequently, no coherent metal melt formed, which sufficiently bonds the metal plate to the cavity support surface and counteracts material distortion. With regard to the other metal plates, which were applied with a lower power PL = 100 W, it can be seen that these also exhibit low material distortion.The optical micrographs show that the metal platelets were bonded to the cavity contact surface with a material bond. For a multiple weld exposure (V), it can be seen that with constant laser power PL = 150 W, there is an increased weld track depth (compare Fig. 7). The metal platelet was heated several times by the multiple weld seam exposure, which significantly increased the molten pool expansion and resulted in sufficient metal platelet adhesion.

7). The metal platelet was heated several times by the multiple weld seam exposure, which significantly increased the molten pool expansion and resulted in sufficient metal platelet adhesion. For the single weld exposure (III), it can be seen that the melt pool expansion was lower and thus the underlying cavity contact surface melted only to a limited extent, as a result of which no sufficiently cohesive metal melt was formed.Fig. 7 Etched weld tracks between metal substrate and full base material1889For the weld width, it can be seen that with an increased weld width of 0.36 mm, a significant ambient gas entrapment (ROI 1) between the metal platelet and the cavity support surface can be detected. This results from the fact that no coherent metal melt of the individual weld tracks was formed and thus individual metal melts lying next to each other were formed.

This results from the fact that no coherent metal melt of the individual weld tracks was formed and thus individual metal melts lying next to each other were formed. For a smaller weld width of 0.18 mm, no significant flaws are discernible because the adjacent molten metal traces formed a coherent metal melt, which also melted the cavity contact surface and thus ensured an expedient metal substrate connection.From the optical micrographs (ROI 2), it can be seen that the metal melt produced in the weld center (test combination 2.3 and 2.5) sinks. This results from the temperature gradients in the molten metal and the associated shear stresses, which led to an asymmetrical molten pool movement (Marangoni convection). Consequently, pressure gradients resulted within the molten metal, which were due to the fact that the molten metal was slowed down at the melting edge. This resulted in a pressure drop at the weld center and a pressure increase at the melt edge.

This resulted in a pressure drop at the weld center and a pressure increase at the melt edge. As a result, the molten metal was conveyed to the melting edge and the molten metal in the weld center sinks. According to Beyer, this lowering of the molten pool describes a budding capillary formation. On the basis of the optical micrographs, the vortex-like molten pool flow caused by the Marangoni effect is evident.In summary, the process parameters investigated (laser power PL = 150 W and scan speed vS = 200 mm/s) with weld exposure V and a weld width of 0.18 mm can be considered suitable for metal substrate attachment of the weldable strain gauge (compare Fig. 8).PL = 150 WTest combination 2.5 Weld seam width = 0.18 mm VvS = 200 mm/sFocus shift zF = 0.00 mm Test combination 2.3 Weld seam width = 0.36 mmFig.

8).PL = 150 WTest combination 2.5 Weld seam width = 0.18 mm VvS = 200 mm/sFocus shift zF = 0.00 mm Test combination 2.3 Weld seam width = 0.36 mmFig. 8 Macroscopic view on the metal plates bonded to the full material IN718 component (width of the metal plate = 10 mm)Implementation and validationIn the following, the elaborated process parameters for the sensor implementation of the strain gauge LS31HT during the PBF-LB/M process were tested in a relevant testing environment.Gas flow measurement componentA gas flow probe (in the following called “probe rake” provided by Vectoflow GmbH was used and produced for the sensor implementation. The said components are usually manufactured by1890IIIthe company Vectoflow using PBF-LB/M to measure locally occurring gas flows in a streaming field. Due to the surrounding flows, vibrations can also occur on the measurement system, which cannot be detected so far.

Due to the surrounding flows, vibrations can also occur on the measurement system, which cannot be detected so far. However, vibration data is an important parameter for the design of the measurement system but also for the even more detailed testing and analysis of flow fields. The following Fig. 9 shows the probe rake manufactured out of IN718 on the SLM125 system from Vectoflow GmbH, with an integrated weldable strain gauge (WSG) using the final parameter constellation (laser power PL = 150 W, scan speed vS = 200 mm/s, weld exposure V, weld width 0.18 mm) for the connection of the metal plate, where the measuring grid of the strain gauge was positioned. For WSG implementation, the cavity was adapted according to the defined cavity layout for the LS31HT strain gauge. The cavity alignment in the X-Y plane resulted from the individual probe rake contour and the associated limited probe rake volume (see Fig. 9).

The cavity alignment in the X-Y plane resulted from the individual probe rake contour and the associated limited probe rake volume (see Fig. 9). Due to the fact that the sensor component requires a cable connection, a cable duct (X-Y plane) was provided, through which the measuring cables of the sensor component are guided out of the probe rake via the carrier plate.Probe rakes on build platformManufactured IN718 probe rakeInterruption plane (X-Y)Position integrated WSG1)2)3)4)5)WSG wiresSupport structuresWSG wires1) Pitot-probe 2) Keel probe (total temperature)3) Keel probe (total pressure)4) 5-hole probe 5) 3-hole probeFig. 9 Additive manufacturing of probe rakes of Vectoflow GmbH with integrated WSGs to determine occurring vibrations and forces.Measurement data evaluationTo evaluate the proper operation of the flow rake, a test of the system for vibration detection by externally applied vibration using MIL STD 810 H was investigated.

The test setup was as shown in Fig. 10, using an MX840B amplifier with a quarter-bridge circuit (350 ohm) at a sampling frequency of 1,200 Hz. A defined frequency of 350 Hz was applied to the flow rake in accordance with MIL STD 810 H via a turbine vibrator. The flow rake was then subjected to a test with a defined frequency of 350 Hz. The correctness of the vibration behavior was ensured by an externally applied Fibre Bragg Grating (FBG) sensor. In this state, the vibrations detected by the1891strain gauge (WSG) were evaluated and compared with those of the FBG. Fig. 10 shows that the power maxima occured equally at 350 Hz.

10 Test setup for determining the correct vibration and spectral power densities of the FBG sensor and the strain gauge integrated in the flow rake in the test setup under an applied vibration of 350 Hz.Summary and OutlookIn this work, a concept was developed with which the strain gauge LS31HT can be inserted into a component made of the metal powder Nickelalloy IN718 during the PBF-LB/M and can be attached to it. It was demonstrated that the design of the weld seams has a major influence on the material bondability of the strain gauge lamina to the rake. After successfully finding a suitable parameter combination, the configuration was used to fabricate a probe rake from Vectoflow. It was then successfully demonstrated that the integrated sensor is capable of detecting relevant vibrations.The investigated manual sensor integration is associated with high costs, since both the personnel and the machine time must be taken into consideration for this.

Furthermore, the achievable accuracy of the integration process depends on the human operator and is therefore not completely reproducible. These disadvantages can be avoided by an automated sensor integration. For this purpose, the already existing cartesian kinematic system (x-y-z-direction) described in [20] should be used for both cavity cleaning and subsequent sensor placement. The kinematic system currently being tested, which can be integrated in the AconityONE PBF-LB/M system, can pick up and place the sensors by means of a vacuum gripper. For this, the cavity must be able to be approached by the kinematic system and accordingly have an opening at the top. The kinematic system, currently under development, has a sufficient placement accuracy for the sensor component. These criteria for cavity design are met by the cavity already described, as are the criteria for sensor placement accuracy by the kinematic system.

These criteria for cavity design are met by the cavity already described, as are the criteria for sensor placement accuracy by the kinematic system. With an automated sensor integration by a kinematic system, the integration time can be reduced to a minimum and the process temperatures as well as the protective gas atmosphere can be kept at a constant level. With the use of automated sensor integration, the problems associated with the conventional method can be avoided and reproducible sensor integration can be ensured.1892375References [1] C. Caviezel, R. Grünwald, S. Ehrenberg-Silies, S. Kind, T. Jetzke, M. Bovenschulte,Aditive Fertigungsverfahren (3D-Druck): Innovationsanalyse, 2017.[2] Wohlers Associates, Wohlers Report 2020: 3D Printing and Additive ManufacturingGlobal State of the Industry, Fort Collins, Colorado, 2020.[3] R. Jiang, R. Kleer, F.T.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIn-Situ Laser Control Method for Polymer Selective Laser Sintering (SLS) Tim Phillips, Austin McElroy, Scott Fish, Joseph BeamanThe University of Texas at Austin Mechanical Engineering Department, Cockrell School of Engineering Austin, TXAbstractSelective laser sintering (SLS) of Nylon is a significant portion of the Additive Manufacturing (AM) market for structurally sensitive applications. To achieve high performance in these laser melted parts, one would like to see consistent melting of the powder over the part region in each layer of the build.

To achieve high performance in these laser melted parts, one would like to see consistent melting of the powder over the part region in each layer of the build. Current research methods into improving this consistency focus on the use of IR sensing to adjust heating elements in attempt to gain even temperature distributions over the pre-laser melted powder layer with the expectation that the laser at constant power, speed, spot size, and spacing will deliver constant melting properties. In this paper we examine a complimentary method of gaining even melted properties by sensing the pre-lased powder along the laser track, and adjusting the laser power to achieve a common post melted temperature everywhere on the part. We describe the feedback based laser control method that varies the laser power in to account for the pre-sintering temperature profile across the part bed.

We describe the feedback based laser control method that varies the laser power in to account for the pre-sintering temperature profile across the part bed. Various tests have been performed, and a method for employing this strategy throughout a build is presented.IntroductionSelective Laser Sintering (SLS) is an additive manufacturing process that creates 3D parts by sintering powdered material, in this case nylon, together with a laser. To create a 3D SLS part, the computer aided design (CAD) model is sliced into a stack of 2D cross-sections, typically 0.004 inches thick each. To begin building a part, the SLS machine uses radiative and conductive heaters to heat the surface of the nylon powder to just below the melting temperature. Once the powder’s temperature has stabilized, the laser will scan the powder in a pattern defined by the first 2D cross-section. Following this, a new layer of powder is spread, heated, and fused in the pattern of the next 2D cross-section.

Following this, a new layer of powder is spread, heated, and fused in the pattern of the next 2D cross-section. This process is repeated until all cross-sections have been sintered and the final 3D part is formed.This process is heavily dependent on the thermal profile of the nylon powder and excess temperature gradients both laterally and in depth can induce curling of the part during the build or inadequate mechanical and dimensional properties [1]. A typical process control objective is to bring the powder to a consistent temperature of around 182ºC for Nylon 12, just a few degrees below the melting point of the powder [2]. Once this temperature has stabilized, the laser scans, using just enough laser power to fully sinter the layer. This ensures that temperature gradients are kept to a minimum. Previous research has examined the design and control of heater systems attempting to maintain small temperature gradients across the powder surface [3] [4].

Previous research has examined the design and control of heater systems attempting to maintain small temperature gradients across the powder surface [3] [4]. These1381systems have reduced, but not eliminated thermal gradients in the pre-scanned powder and further improvement is needed to provide consistent/predictable part properties.Other factors that affect the thermal profile of the powder are part geometry and laser scan pattern. While nylon powder is a good thermal insulator, heat still dissipates between layers. This means that in instances where the current layer being sintered has regions beneath it that vary in amount of sintering, such as when creating an unsupported overhang, thermal gradients can arise.The most basic laser scan pattern consists of straight lines that overlap slightly along their axis, meaning that the first few lines have less overlap with previously sintered lines when compared to lines in the center of the part.

This can be seen in Figure 1, where three rectangular part sections are being scanned with a fill pattern, and the first 3 vector scans are cooler than the later 8 scan lines. This figure also shows hot spots at the beginning and end of the scan lines due to inadequate galvanometer tuning.Figure 1: Laser Scan Line TemperatureIn an open loop system, these influences on powder temperature are unaccounted for and can decrease the quality of the final part. When a constant laser power is used along a scan line, the sintered powder will largely all experience the same temperature increase (disregarding end of vector over-sintering caused by inadequate galvanometer tuning). This means that any temperature gradient that exists in the powder initially will be preserved after sintering. The resulting part also contains this temperature gradient which can cause the part to curl or can cause unpredicted geometrical and structural characteristics.

The resulting part also contains this temperature gradient which can cause the part to curl or can cause unpredicted geometrical and structural characteristics. By controlling the post-sintering temperature of a part and decreasing the variation among post-sintering temperatures, part quality and predictability can be increased [5]. Post-sintering temperature monitoring [6] and control [7] has been explored previously for metal additive manufacturing processes, but little has been done to control post-sintering temperatures in polymer SLS.1382Experimental SetupTestbedIn-situ laser power compensation using closed-loop feedback was accomplished on the Laser Additive Manufacturing Pilot System (LAMPS), seen in Figure 2. LAMPS is an open hardware SLS platform that was designed and built at the University of Texas at Austin as an experimental testbed for SLS process control. LAMPS uses a 60 watt CO2 laser with a Cambridge Technologies EC1000 laser and galvanometer controller.

LAMPS uses a 60 watt CO2 laser with a Cambridge Technologies EC1000 laser and galvanometer controller. The build surface is 220 mm x 220 mm and sintering takes place in a heated nitrogen atmosphere. The machine utilizes independently controlled strip heaters to heat the walls of the build chamber which, in turn, control the atmosphere temperature. The strip heaters do the majority of the work heating the powder and additional quartz lamps are used to fine tune the temperature as well as to do the fast heating required after new powder is spread. All measurement and control is handled through National Instruments Labview software and cRIO hardware.Figure 2: LAMPS renderingInfrared Camera CapabilityIn addition to atmosphere feedback via thermocouples, LAMPS uses two infrared (IR) cameras to monitor and control the building process. A FLIR A325 Long Wave Infrared (LWIR) camera is held in a stationary position and directed at the build surface.

A FLIR A325 Long Wave Infrared (LWIR) camera is held in a stationary position and directed at the build surface. This camera provides feedback to control the duty cycle of the quartz lamps. This allows for control of the powder surface temperature as each layer is spread in preparation for laser melting. The second IR camera is a FLIR SC8240 Mid Wave Infrared (MWIR) camera. This is a high-speed IR camera capable of recording 2,243 frames per second at 64x64 pixels per frame. This camera is boresighted with the CO2 laser, providing a constant view of the build surface in the immediate1383vicinity of the scanning laser contact point. This is accomplished by placing a dichroic mirror in the laser path before the galvanometer system, as seen in Figure 3. Since this is done prior to the optical path entering the galvanometer system, the field of view (FOV) of the MWIR camera will not be stationary and will follow the laser as the galvanometer scans across the powder surface.

The result is a close up view of the powder surface with the laser spot centered in the image, regardless of position on the build surface.An image taken with the MWIR camera is presented in Figure 4. The color bar shown in the image is in units of “count”, which is a measure of the amount of photons reaching the camera’s sensor during the integration period. It is possible to convert this unit to temperature, but it is highly dependent on parameters such as transmittance of the optics, emissivity of the powder, reflected radiation, and other parameters that are not known precisely. Since counts can be related to temperature however, and our interest is in creating consistency and temperature, the counts can be used directly in the feedback control.

Since counts can be related to temperature however, and our interest is in creating consistency and temperature, the counts can be used directly in the feedback control. To give the reader a sense of scale, 7900 counts is roughly equal to 177ºC and a 400 count change is equal to a change of roughly 3.8ºC.Figure 3: Boresight IR Camera Setup1384Figure 4: MWIR Camera ViewControl StrategyThe results presented in this paper employ the strategy described here to create individual laser scan lines. By repeating this process, multiple scan lines can be positioned and overlapped, creating an entire 3D part with excellent control of the post-sintering temperature of the part. This in-situ laser compensation strategy employs a multiple scan approach that is separated into two phases. In the first phase, the galvanometers scan the MWIR camera and laser across the build surface in the same fashion as if it were sintering a layer, except with the laser turned off.

In the first phase, the galvanometers scan the MWIR camera and laser across the build surface in the same fashion as if it were sintering a layer, except with the laser turned off. During this phase, the MWIR camera is used to record powder temperature where the laser will sinter, giving the initial temperature gradient of the powder. This temperature gradient is analyzed and used to develop a laser power profile for sintering that will produce a scan line with a consistent post-sintering temperature, regardless of the temperature gradient present initially.The second phase is to employ the scan strategy developed in the first phase to sinter the scan line with variable power. The MWIR camera continues to record data in this phase and is used to verify the post-sintering temperatures.Hardware LimitationsThe ultimate goal of this technology is to implement a real-time control method, where the system can acquire feedback and vary the laser power in real time during a laser scan.

Current hardware compatibility limitations, described below, prohibit real-time control for this experiment so results presented in this paper focus on in-situ control. The compatibility issues will be addressed in future effort and publications.1385One current limitation to using a variable laser power is with the laser controller. LAMPS uses a Cambridge Technologies EC1000 controller for the laser and galvanometer system. The EC1000 only allows a fixed power to be used for each laser scan line, meaning a line must be broken into subsections and each fed separately to the EC1000 in order to use different amounts of laser power in different areas of the original scan line. Another limitation is the framerate of the MWIR camera.

Another limitation is the framerate of the MWIR camera. While it records at a high framerate of 2,243 frames per second, a framerate greater than 4,000 frames per second would be needed to sample each voxel of powder [8].As mentioned previously, the machine control is implemented using LabVIEW and incorporating the MWIR camera into the LAMPS control program was not completed for this experiment. The host computer is not fast enough to handle the control program simultaneously with the high data rate MWIR data acquisition. The interim solution was to use a separate computer for recording and analyzing the MWIR data and manually transferring the scan strategy to the LAMPS control computer to actuate the laser. This resulted in a delay of approximately two minutes between acquiring the initial temperature gradient and actuating the laser, hence the inability to do real-time control.

This resulted in a delay of approximately two minutes between acquiring the initial temperature gradient and actuating the laser, hence the inability to do real-time control. This is also the reason the results here are limited to individual scan lines, as doing multiple lines would change the temperature gradient and phase one of the control strategy would have to be repeated before every scan line. By incorporating the MWIR camera and LAMPS control software on the same computer, though, it would be possible to use this control strategy to build an entire 3D part. The objective here is to show feasibility and benefits of the general approach, even with these current limitations.ResultsBaseline Fixed-Power ResultsOne assumption made was that using a fixed laser power would preserve the initial temperature gradient. To verify this assumption, the MWIR camera was scanned across a laser scan line before and during sintering to record temperatures of where the laser sintered.

To verify this assumption, the MWIR camera was scanned across a laser scan line before and during sintering to record temperatures of where the laser sintered. Figure 5 shows these results for scanning a line with 10% laser power. The green line is the temperature before sintering and the blue line is the temperature after. To better visualize how the post- sintering temperature is affected by a constant laser power, the second subplot shifts the pre- sintering temperature upwards by a constant values so that the beginning of the scan lines coincide. This clearly shows that the temperature profile is the same for both lines, meaning it was unaffected by the constant laser power sintering. Figure 6 shows the results of a similar test, this time with a scan line split into two subsections with each subsection receiving a different laser power.

Figure 6 shows the results of a similar test, this time with a scan line split into two subsections with each subsection receiving a different laser power. Again, this shows that the pre-sintering temperature profile has been preserved and the assumption is valid.1386Figure 5: Temperature Comparison at 10% Laser PowerFigure 6: Temperature Comparison with Variable Laser Power1387To accurately control the post-sintering temperature, a transfer function between applied laser power and post-sintering temperature increase must be created. To do this, multiple lines were sintered using different laser powers and their pre- and post-sintering temperatures were compared. The result of one of these tests is seen in Figure 7. Once these tests were performed, the results were compiled and a best-fit line was found that would provide the equation for desired temperature increase and delivered laser power.

Once these tests were performed, the results were compiled and a best-fit line was found that would provide the equation for desired temperature increase and delivered laser power. This can be seen in Figure 8, where some outliers in the data were removed and the resulting 4th order polynomial fit line has an R2 value of 0.9988. Note that this figure shows 5% laser power results in no temperature increase. This is because that is near the threshold value for laser power, below which the laser does not fire.Figure 7: Laser Power and Temperature Increase Trial1388Figure 8: Temperature Increase vs. Laser PowerSome baseline testing was conducted using a fixed laser power, the results of which will be compared with the results of the in-situ control method presented.

Laser PowerSome baseline testing was conducted using a fixed laser power, the results of which will be compared with the results of the in-situ control method presented. The raw data is shown in Figure 9 which shows a dwell time at the beginning of the test before the galvanometers begin moving, a temperature rise as the galvanometers move to the start of the scan line immediately followed by a temperature decrease as the galvanometers scan across the line with the laser off. This is immediately followed by a temperature decrease as the galvanometer moves back to the start of the scan line then a large jump in temperature as the laser is turned on and the line is scanned. This is followed by another dwell time as the galvanometer rests at the end of the scan line with the laser turned off and the powder can be seen cooling.

This is followed by another dwell time as the galvanometer rests at the end of the scan line with the laser turned off and the powder can be seen cooling. Trimming this data to just show the two sections where the galvanometers are scanning forwards across the scan line (once with the laser off, once with it on) is seen in Figure 10 where the data is shifted in the x-direction such that the pre- and post-sintering temperatures are lined up with their position in the build chamber. The results of the three baseline tests are quantified in Table 1.1389Figure 9: Raw Baseline Temperature DataFigure 10: Baseline Temperature Data1390Table 1: Baseline Temperature Measurements, Measured in Photon CountTrial Avg. Pre- Sintering Temp.1 2 38022.6 7843.7 7751.7Max Pre- Sintering Temp Difference 291 314 316Pre- Sintering Temp STD 88.6 93.2 97.8Avg.

Pre- Sintering Temp.1 2 38022.6 7843.7 7751.7Max Pre- Sintering Temp Difference 291 314 316Pre- Sintering Temp STD 88.6 93.2 97.8Avg. Post- Sintering Temp 9584.1 9442.4 9292.1Max Post- Sintering Temp Difference 522 256 414Post- Sintering Temp STD 89.5 73.2 75.8Temp Diff Change79.38% 18.47% -31.01%Temp STD Change1.02% 21.46% 22.49%In-Situ Laser Power CompensationIn-situ control of laser power was performed as described in the control strategy section. The result of phase one of the control strategy, where the pre-sintering temperature profile is acquired, for one of the in-situ control trials can be seen in Figure 11. With a goal of a post- sintering temperature of 9,000 counts, this temperature profile was fed through the laser power to temperature increase transfer function. The result of the transfer function is seen in Figure 12. Each time the laser power is changed a new line subsection must be created.

The result of the transfer function is seen in Figure 12. Each time the laser power is changed a new line subsection must be created. To keep the number of line subsections manageable, the laser power percentage was limited to integer values. If the laser controller was capable of modifying laser power mid-scan line, this would look like a smooth function rather than a step-wise function.Figure 11: Pre-Sintering Temperature Profile1391Figure 12: Delivered Laser PowerThe scan strategy detailed in Figure 12 was implemented and the results can be seen in Figure 13. In this image, the red line represents the same data seen in Figure 11 and is the temperature profile that is used to compute the laser power profile. Once the laser power profile is developed, it is fed back through the laser-power-to-temperature-increase transfer function to get a projected temperature increase.

Once the laser power profile is developed, it is fed back through the laser-power-to-temperature-increase transfer function to get a projected temperature increase. This temperature increase is added to the initial temperature profile to get a projected post-sintering temperature profile, which is shown in black. Note that this is not a straight line at 9,000 counts because only integer laser power settings were used as discussed above.The process of analyzing the initial temperature profile, developing a laser power profile, and preparing the laser to implement said power profile takes approximately two minutes due to the restrictions detailed in the hardware limitations section. Immediately prior to the variable laser power scan line, the galvanometers scan the line again with no laser power. The result of this is shown in blue.

Immediately prior to the variable laser power scan line, the galvanometers scan the line again with no laser power. The result of this is shown in blue. The purpose of this step is to ensure that the initial temperature of the scan line has not changed in the two minutes it took since the initial data was taken. If the red and blue lines are significantly different, the system has changed and it is expected that the laser power profile developed would not produce the desired results. In this case, the red and blue lines are in close agreement, so the actual post-sintering temperature, shown in green, is expected to match closely with the projected post-sintering temperature shown in black. This figure shows a successful trial as the post-sintering temperature was well controlled and is not influenced by the initial temperature profile of the powder.1392Figure 13: In-Situ Laser Control Results for Trial 4Not all trials were as successful, as can be seen in Figure 14.

This trial shows overcompensation by the laser as the post-sintering temperature increases over the length of the scan line. One explanation for this result is that this trial was performed at the lowest power level and it is possible that the laser power to temperature increase transfer function was not as accurate in this region. The results of all closed-loop trials are quantified in Table 2. Note that this table is missing data for trials 3, 8, and 13, which contained MWIR camera recording errors or errors found after the test in one of the MATLAB scripts used to create the variable laser power scan files. These are all errors in the implementation of control, not flaws in the control method and do not detract from the validity of the in-situ laser compensation strategy presented.

This revealed the need for an improved laser control method that could compensate for initial temperature gradients and produce sintered regions with little temperature variation. The method proposed used an infrared camera to measure the pre-sintering powder temperature and compute an appropriate laser power needed to produce the desired post-sintering temperature. This regulated the laser power delivered in order to counteract the initial thermal profile of the powder to produce a uniform post-sintering temperature.The results are overwhelmingly positive, with each trial outperforming the baseline open loop system. The control strategy presented in this paper saw reductions in the influence of the pre-sintering temperature profile of up to 65%.

The control strategy presented in this paper saw reductions in the influence of the pre-sintering temperature profile of up to 65%. This increased control over post-sintering temperatures has the capability of improving the mechanical and dimensional properties of SLS parts [5], as well as increasing the repeatability of the sintering process.Future work involves applying the control strategy presented to an entire 3D part using real-time control. This will involve further automation in developing the laser power profile for each scan line. It would also be beneficial to have an automated process for developing the laser power to temperature increase transfer function. This process could be performed periodically to ensure the delivered laser power is still expected to produce the same temperature increase. Periodically updating the transfer function is likely to increase the amount of control over the post-sintering temperature.

Periodically updating the transfer function is likely to increase the amount of control over the post-sintering temperature. Another avenue for further exploration is to adapt this in-situ control strategy into a real-time control strategy. This will likely involve additional hardware and software that operates quickly and is capable of varying the laser power in real-time without the need to split the scan line into subsections.1395Bibliography[1] J. A. Benda, "Temperature-Controlled Selective Laser Sintering," in Solid Freeform FabricationSymposium, Austin, 1994.[2] P. Hall, Interviewee, SLS Process Overview. [Interview]. 2015.[3] D. L. Bourell, T. J. Watt, D. K. Leigh and B. Fulcher, "Performance Limitations in Polymer LaserSintering," Physics Procedia, pp. 147-156, 2014.[4] W. W. Wroe, "Improvements and Effects of Thermal History on Mechanical Properties for PolymerSelective Laser Sintering (SLS)," University of Texas, Austin, Austin, 2015.[5] J. R.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIN-SITU LAYER-WISE QUALITY MONITORING FOR LASER-BASED ADDITIVE MANUFACTURING USING IMAGE SERIES ANALYSISMehrnaz Noroozi Esfahani*, Linkan Bian*†, Wenmeng Tian*†Department of Industrial and Systems Engineering, Mississippi State University, Mississippi State, MS 39762 †Center for Advanced Vehicular Systems (CAVS), Mississippi State University, MS 39762AbstractQuality assurance has been one of the major challenges in laser-based additive manufacturing (AM) processes. This study proposes a novel process modeling methodology for layer-wise in-situ quality monitoring based on image series analysis. An image-based autoregressive (AR) model has been proposed based on the image registration function between consecutively observed thermal images.

An image-based autoregressive (AR) model has been proposed based on the image registration function between consecutively observed thermal images. Image registration is used to extract melt pool location and orientation change between consecutive images, which contains sensing stability information. Subsequently, a Gaussian process model is used to characterize the spatial correlation within the error matrix. Finally, the extracted features from the aforementioned processes are jointly used for layer-wise quality monitoring. A case study of a thin wall fabrication by a Directed Laser Deposition (DLD) process is used to demonstrate the effectiveness of the proposed methodology.1. IntroductionAdditive Manufacturing (AM) processes add material in layer-by-layer fashion to achieve the final geometry, which enables new design options that cannot be obtained through conventional manufacturing technologies [1].

Nevertheless, quality issues of AM parts are major barriers preventing wider industrial adoption of AM. Due to the existence of defects such as porosity and cracking, the mechanical properties of additively manufactured parts can hardly satisfy the strict requirements of industrial application. AM processes are governed by the process-structure- property (PSP) relationship. Once the PSP relationship is established, the build structure and properties can be predicted. The melt pool images are regarded as a most informative process signature for build structure prediction. In this paper, a novel layer-based quality characterization method is proposed for layer-wise anomaly detection for potential process correction/control actions during the fabrication.

In this paper, a novel layer-based quality characterization method is proposed for layer-wise anomaly detection for potential process correction/control actions during the fabrication. Figure 1 shows a schematic plot of a Direct Laser Deposition Process (DLD) fabrication process monitored by a coaxial pyrometer.The challenges confronting layer-based process anomaly detection are two-fold: 1) there is a huge amount of uncertainty in the relationship between the compositional and process parameters in the basic thermo-mechanical process of metal printing; 2) the advanced sensing technologies provide high volume of noisy thermal history data that represent complex spatio-temporal relationship. To model correlation between thermal history and microstructure properties, finite element methods (FEMs) have been widely used.

The thermal behavior and temperature distribution and their effects on the stress, formation, and hardness of parts along with phase transformation during the AM process have been thoroughly studied in the literature [2, 3].202However, there are some challenges in the FEM based approaches that hinder their application in online prediction [4]: 1) high computational cost required, making it extremely time consuming to implement in real time; 2) their high dependency on part geometry which makes it almost impossible to generalize between different designs; and 3) the deterministic nature of FEM related approaches, making it difficult to incorporate all sources of process uncertainty in the modeling for anomaly detection.Laser SourcePyrometerProne to porosityL Abnormal rlaycr heightTimeFigure 1: DLD process monitored by a co-axial pyrometerThe advancement in sensing technologies has enabled real-time monitoring of thermal history for anomaly detection via infrared thermography, generating thermal imaging data streams with extremely large volume and complex structure.

Although existing data-driven methods consider robust statistical methods to identify anomalies from thermal image [5-6], local features for prediction purposes are used in most of the existing work and, consequently, they cannot be directly utilized to distinguish the profile of an entire deposited layer. This being the case, layer- wise process signature has received the attention of the AM community. Liu et al. [7] has modeled layer-wise spatial porosity evolution based on X-ray computed tomography (XCT), which is an expensive and time-consuming process. Furthermore, optical imaging systems have been used in laser powder bed fusion systems for layer-wise monitoring [8].

Furthermore, optical imaging systems have been used in laser powder bed fusion systems for layer-wise monitoring [8]. Moreover, layer-wise anomaly detection based on thermal history has also been proposed which incorporates a tensor-based dimension reduction and a convex hull based variability characterization approach [9].In this paper, a novel data-driven methodology is proposed to extract layer-wise process signatures from real-time thermal history for direct laser deposition (DLD) processes. AM process knowledge is leveraged to develop the novel approach. Therefore, the major assumption of the proposed methodology is that a stable thermal history leads to homogenous microstructure and thus porosity-free deposited layers. As illustrated in Figure 1, when there is a layer height anomaly occurring in the process, the observed melt pool in the thermal image will shift in both shape and location.

As illustrated in Figure 1, when there is a layer height anomaly occurring in the process, the observed melt pool in the thermal image will shift in both shape and location. Based on this observation, process features can be proposed to quantify the change in the thermal image series obtained from fabricating one layer. Therefore, process features can be extracted from two sources in the thermal image series, 1) image registration information; and 2)203first order difference of the registered thermal images. Subsequently, the extracted features are used for anomaly detection for each deposited layer in the DLD process.The rest of the paper is structured as follows. Section 2 provides the proposed framework for layer-wise process feature extraction and anomaly detection, while Section 3 uses a case study of fabricating a thin wall using a DLD process to validate the proposed methodology. Finally, Section 4 summarizes the conclusion and potential future research directions.2.

Finally, Section 4 summarizes the conclusion and potential future research directions.2. Methodology2.1.Data Description and ChallengesThermal history indicates the thermal response during AM fabrication as a function of time, which is represented by an image series (video) captured by a pyrometer camera during the build. In the thermal images, a region of superheated molten metal is defined as the melt pool [10]. As a key signature of the fabrication process, the melt pool initiates the solidification process in AM and thus is informative to predict part porosity information. Each melt pool image contains location-based temperatures, and each layer can be considered as a series of images. Most of the currently available approaches use individual melt pool image for anomaly detection at the specific location where the melt pool is observed during the fabrication.

Most of the currently available approaches use individual melt pool image for anomaly detection at the specific location where the melt pool is observed during the fabrication. However, plenty of spatio- temporal correlation information presents in the image series which can be used for anomaly detection and quality prediction. The monitoring of layer-wise thermal history is challenging due to 1) complex spatio-temporal correlation, 2) high dimensionality of thermal images, 3) discrete data sampling, and 4) unreliable and missing data.2.2.Thermal Image Series AnalysisIn this subsection, the proposed framework for modeling the thermal image series collected from fabricating one layer is briefly described.

(cid:2013)(cid:3047)(cid:2878)(cid:2869) denotes an error matrix (with a dimension of (cid:1835) × (cid:1836)) representing the piece of information that cannot be explained by the registration operation. A Gaussian process model is used here to characterize the error term (cid:2013)(cid:3047)(cid:2878)(cid:2869), and the detailed information is included in subsection 2.2.2. Figure 2 shows the consecutive melt pool images as an image series, and briefly illustrates the proposed formulation in Equation (1).204tt+lt+2t+3TimeFigure 2: Layer-wise thermal image series formulation2.2.1 Image RegistrationImage registration is a widely used image processing technique that reveals some hidden relationship between the input and the reference images in applications such as health care industry [11], computer vision for target localization, automatic quality control, motion tracking, and cartography for map updating.

It is worth noting that Quasi-Newton optimizer is used for parameter estimation in the GP.jj'i'Figure 3: Distance between pixels in a thermal image2.3.Layer-wise Feature ExtractionThere are two groups of features extracted to characterize the thermal image series: one is extracted from the transformation matrix obtained from the registration function, and the other is extracted from the estimated parameters from the GP models to characterize the error matrix.206(3)Image registration related features are extracted from the transformation matrix (cid:2743)(cid:3047).

Supervised learning methods can be used to train a classification model. Support Vector Machine (SVM) is considered as a powerful classification technique with a diversity of kernel functions. By finding the best hyperplane that separates all data points of different classes, an SVM classifier can be trained using data labeled with X-ray CT evaluations. More specifically, a layer that contains at least one porosity is labeled as unhealthy, and a porosity-free layer is labeled as healthy. Among SVM classifiers, linear SVM is applied in this paper for simplicity and better interpretability. To achieve more flexibility, more sophisticated kernels such as Gaussian kernel or polynomial kernel can be used [15]. It is worth noting that various other machine learning algorithms can also be applied to this problem. Comparing classification performance of different machine learning schemes is out of the scope of this paper.3.

Comparing classification performance of different machine learning schemes is out of the scope of this paper.3. Case StudyThe performance of the proposed methodology is validated using a direct laser deposition process. A thin wall of 60 layers in total was fabricated using Ti-6AL-4V. During the fabrication, the thermal images of the melt pools related to different locations on the thin wall are captured by a pyrometer camera. The number of thermal images captured using co-axial pyrometer camera is 1564 with each represented by a 480 × 752 matrix of pixels. The process setup parameters are summarized in Table 1 [5]. After the build was completed, the porosity structure inside the build was identified using an X-ray CT scan (Skyscan 1172). The X-ray examination indicated that 26 layers out of 60 layers of the thin wall include at least one porosity. Figure 5 illustrates a melt pool image series within a layer containing a porosity related to an unhealthy melt pool at t+17.

Figure 5 illustrates a melt pool image series within a layer containing a porosity related to an unhealthy melt pool at t+17. It is worth noting that minimum size for a pore is 0.05(cid:2020)(cid:1865).Table 1: LENS process design parameters for the thin wall.Scan speed12.70 mm/sStarting offset from substrate130.391mmPowder feed rate0.060 g/sDetermination of layer thickness0.508 mmDetermination of hatch spacing0.508 mm Nozzle diameter0.889 mmPower300 WSubstrate (stainless steel) 3.175 mm208t+lt+2t+3t+17t+20Figure 5: The image of a thin wall containing one pore in an unhealthy layerFor image registration, due to the printing path design was unidirectional, the orientation of the melt pool image should not change over time. Therefore, only translation operation is estimated in the transformation matrix (cid:2743)(cid:3047). Fabricating a thin wall, the melt pool only shifts from image to image and the melt pool is not subjected to rotation.

Fabricating a thin wall, the melt pool only shifts from image to image and the melt pool is not subjected to rotation. Regarding that the only transformation that can be applicable is translation, the feature extracted from translation are the only feature extracted from the image registration process. The error term (cid:2013)(cid:3047)’s were modeled using Gaussian process models, and the layer-wise features were obtained by calculating the maximum values for each feature.Leave-one-out validation was used to test the effectiveness of the proposed method. Table 2 illustrates the confusion matrix after implementing the aforementioned model. Four layers out of 60 layers (6.67% of the total layers) are misclassified, in which two misclassifications for healthy and two for unhealthy.

TP shows true positive layers which are healthy and predicted accurately as unhealthy. TN stands for true negative that illustrated the items that are accurately predicted as healthy. On the other hand, FN defines inaccurate prediction that are actually unhealthy, while FP shows inaccurate prediction that are actually healthy.209Table 2: Confusion Matrix for Leave-one-out Cross ValidationPredicted Healthy UnhealthyIActualHealthy UnhealthyII32 2I2 24The results indicate that the recall and precision of leave-one-out cross validation are 0.9231 and 0.9231, respectively, and consequently, the F-score is also 0.9231. This indicates that the classification model accuracy is reasonably high and can be used for layer-wise anomaly detection. One of the possible reasons for misclassification is the discrete data sampling for thermal history fails to capture the thermal image when the anomaly occurs. Consequently, some information will be missed.

Consequently, some information will be missed. In addition, the X-ray CT scans may be subject to noise and error as well. Moreover, the proposed anomaly detection method does not consider effects of re-melting, which can potentially correct some of the porosities generated in the previous layer.4. ConclusionLack of repeatability in the additively fabricated parts is one of the major challenges that hinders broader industrial applications of additive manufacturing processes. Possible build quality issues, such as internal porosity, deformation, and cracking, can lead to significantly compromised mechanical properties. Comprehensive studies including data-driven methods have been focused on local characterization for anomaly detection based on single thermal images. A layer-wise thermal image modeling framework can take into account the complex spatio-temporal relationship within the image series, which can potential achieve improved anomaly detection results.

A layer-wise thermal image modeling framework can take into account the complex spatio-temporal relationship within the image series, which can potential achieve improved anomaly detection results. In this paper, the layer-wise thermal images are formulated as an image series, and an image-based autoregressive (AR) model has been proposed based on the registration function between consecutively observed thermal images. Multiple features are extracted from the AR model for each consecutive pair of images, and layer-wise features are extracted subsequently. Support Vector Machine method is used for anomaly detection based on the extracted layer-wise features. A case study based on a thin wall fabrication using a DLD process is used to validate the proposed methodology. The classification accuracy of the proposed method is reasonably high which makes the model capable of predicting defects.

The classification accuracy of the proposed method is reasonably high which makes the model capable of predicting defects. For future works, the represented model can be applied to the thermal image analysis of AM parts with more complicated geometry.5. AcknowledgmentResearch was sponsored by the Army Research Laboratory and was accomplished under Cooperative Agreement Number W911NF-15-2-0025. The views and conclusions contained in this document is those of the authors and should not be interpreted as representing the official policies, either expressed or implied, of the Army Research Laboratory or the U.S. Government. The U.S. Government is authorized to reproduce and distribute reprints for Government purposes notwithstanding any copyright notation herein.6. References210[1] M. Grasso and B. M. Colosimo, “Process defects and in situ monitoring methods in metal powder bed fusion: a review,” Measurement Science Technology, vol. 28, no. 4, p. 44005, Apr. 2017.[2] M.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceIn-situ Local Part Qualification of SLM 304L Stainless Steel through Voxel Based Processing of SWIR Imaging DataCody S. Lough1, Xin Wang1, Robert G. Landers1, Douglas A. Bristow1, James A. Drallmeier1, Edward C. Kinzel21Department of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 654092Department of Aerospace and Mechanical Engineering, University of Notre Dame, Notre Dame, IN 46556AbstractThis paper demonstrates the potential for qualification through local part property prediction of 304L stainless steel parts manufactured by Selective Laser Melting (SLM). This is accomplished through voxel based processing of SWIR imaging data measured in-situ.

This is accomplished through voxel based processing of SWIR imaging data measured in-situ. Thermal features are extracted from time-series SWIR imaging data recorded from layer-to-layer to generate 3D point cloud reconstructions of parts. The voxel based data is indexed with localized measurements of SLM part properties (light-to-dark microstructural feature ratio, microhardness, μCT data) to demonstrate the correlations. Various features are extracted from the thermal history for comparison of their respective abilities to predict the resulting local part properties. The correlations and comparisons developed in this paper are then used to discuss the capability of a voxel based framework using information from in-situ measurements of the thermal history to locally qualify parts manufactured by SLM.

This work was funded by the Department of Energy’s Kansas City National Security Campus which is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1.0. IntroductionThe incorporation of in-situ radiometry during Selective Laser Melting (SLM), a powder bed based Additive Manufacturing (AM) technology, provides a better understanding of the manufacturing process through non-contact measurement of emitted signals. The measurement and processing of radiometric signals provides insight into the complex thermal history experienced by a part during the manufacturing process. The complex thermal history is a result of the laser processing parameters, scan pathing, and part geometry, and can vary significantly in the volume of parts. This is an issue since the thermal history (along with material chemistry) determines the part microstructure and mechanical properties.

This is an issue since the thermal history (along with material chemistry) determines the part microstructure and mechanical properties. The variances in the thermal history can lead to non-uniformity in the properties of SLM parts and defects. Measuring the variances in the thermal history allows informed decisions to be made in the production of SLM parts. The information from radiometric measurement methods can be used to identify defects, develop correlations with part properties, and make predictions of part properties resulting from SLM.Various instrumentation including visible cameras, infrared cameras, pyrometers, and photodiodes have been used in studies to obtain information (apparent melt pool geometry, intensity/temperature, laser spatter, etc.) from the radiometric signals [1].

The features extracted from the measurements performed during the SLM process have been analyzed for various1611processing conditions [2] and mapped to generate 2D and 3D data sets for parts [3]. Data processing methods have been developed for analysis of both time series data and voxel based maps of powder bed based AM parts to identify defects (melt pool quality, variances in the thermal history, porosity, failure of parts, etc.) [4-10]. Efforts are found in the literature to explore how these defects change with SLM processing conditions with an example in [11]. The understanding of this relationship has been expanded in a limited number of reports by correlating the AM part properties (tensile strength, porosity) with the information extracted from radiometric measurements [12-14].The studies conducted to understand how radiometric signals relate to part quality and properties have shown the technique to be success.

However, there is an ongoing need to mature the radiometric monitoring methods for the SLM process to enable in-situ part qualification with high confidence and simultaneous prediction of the part properties. The ability to predict bulk part properties has been explored through the development of models informed by FEA, or experimental part property measurements. In these works, the predictions of SLM part quality and properties include residual stress [15,16], porosity [17,18], mechanical performance [18], surface roughness [19], and microstructure [20]. These bulk predictions of part properties are valuable, but understanding the thermal history through experimental radiometric measurements permits the ability to predict part properties on a local basis. The radiometric measurements are necessary due to the previously discussed complex nature of the SLM process, where the variances in the local thermal history can affect the properties.

The radiometric measurements are necessary due to the previously discussed complex nature of the SLM process, where the variances in the local thermal history can affect the properties. The ability to make predictions of the variances in local part properties adds value to a radiometric process monitoring framework by providing a fundamental insight into how measured local processing conditions ultimately relate to the SLM part properties.There are very few reports in the literature in the high potential area of predicting local SLM part properties through in-situ radiometric measurements. Coeck et al. demonstrated that melt pool event information measured with photodiodes can be used to detect and predict porosity in samples with high efficiency [21]. While this approach works well, recording the entire spatial and temporal components of the thermal history through thermographic imaging could provide some advantages over photodiodes with the cost of sampling frequency.

Recording the spatial and temporal components of the SLM thermal history permits the effects of material re-melting to be included in predictions of part properties. Data processing of measurements from thermographic imaging can also be expanded to incorporate spatial (temperature gradients) and temporal derivatives (cooling rates), which are related to the final part microstructure.In this paper, the ability to use thermographic imaging in the prediction of local SLM part properties (microstructure, microhardness, porosity) is explored as a part of an effort toward local part qualification. The advantage of this method to account for re-melting of material was used in this work (incorporation of derivative data will be completed in the future). To accomplish the goal of local part property predictions, a Short-Wave Infrared (SWIR) imaging camera was used to record the thermal history during SLM of 304L stainless steel.

To accomplish the goal of local part property predictions, a Short-Wave Infrared (SWIR) imaging camera was used to record the thermal history during SLM of 304L stainless steel. Thermal features were extracted from the time-series SWIR imaging data (time above threshold and maximum radiance) and used to develop models for predicting the local part properties. The framework was then used to predict the local part properties of an SLM part with a complex embedded feature.16122.0. Experimental Setup2.1. Samples for Local Property Correlations and PredictionsThe two samples illustrated in Fig. 1 were manufactured by using a Renishaw AM250 and 304L stainless steel powder. The AM250 is a SLM system that processes samples with a fiber laser (Pmax = 200 W, (cid:540) = 1070 nm) scanning with a point-to-point exposure strategy. The first sample was used to develop correlations between thermal features measured through SWIR imaging and local SLM part properties. This sample (Fig. 1.

The first sample was used to develop correlations between thermal features measured through SWIR imaging and local SLM part properties. This sample (Fig. 1. (a)) was partitioned into 20 sections of 20 layers for assigning various laser processing parameters. This generated a range in local thermal histories and part properties. The laser processing parameters varied in this work included the power, P = 100 – 200 W, and exposure time, te = 50 – 125 μs (point distance, dp = 60 μm, and hatch spacing, sh = 85 μm, held constant). The various process parameter combinations resulted in fluence values (optical energy per unit area) ranging from 1.3 J/mm2 to 6.5 J/mm2. The various fluence values were randomized before assigning them to the sections of the part. Manufacturing the first sample by this procedure assumed that an increasing the distance in the build direction does not significantly impact the thermal history of parts for a constant layer-to-layer time.

The procedure also assumes layer-to-layer steady state is reached for the part thermal history in the base (50 layers) and changing the parameters between the sections resulted in convergence to steady state for that parameter set immediately.(a) 1 11 2222 2 22 3C= @!1i) ~ ~(b)n o i t c e S(cid:194) (cid:194) (cid:194)GJ l[ IBuild direction(M) ~ (;2! ~202 mmFig. 1. CAD model of (a) various parameters sample and (b) embedded feature sample cross- section.The correlations developed by analysis of the various parameter set sample were used with new SWIR imaging measurements from a second sample to predict its local part properties. This second SLM 304L sample (Fig. 1. (b)) was manufactured with an embedded feature of text processed with a laser power of 100 W (surrounding geometry processed with 200 W). Both the surrounding geometry and text features of this second sample were built with border scans.

Both the surrounding geometry and text features of this second sample were built with border scans. Processing the second sample in the described way led to the potential for local variances in part microstructure and mechanical properties due to several factors. The first factor was the lower laser power leading to sub-optimal mechanical properties and an increase in lack of fusion porosity in locations corresponding to the text interior. Additionally, at the edges of the text border scans1613re-melting material could cause differences in microstructure due to changes in thermal boundary conditions during and after solidification. The border scans could have also led to an increase in porosity due to keyholing (repeated energy input in localized region). Finally, the differences in cross-sectional area processed for the text and surrounding regions led to changes in scan pathing from layer-to-layer. This introduced the potential for geometric dependence for local part properties.

This introduced the potential for geometric dependence for local part properties. The predictions of the local part properties for the embedded feature sample were validated through experimental measurements.2.2. Voxel based Processing of SWIR Imaging DataThe Renishaw AM250 used to manufacture the samples for this work contains a FLIR SC6201 SWIR imaging camera (sensitive: (cid:540) = 0.9-1.7 μm, filtered: (cid:540) = 1.45±0.05 μm) installed in a staring configuration to record the SLM process layer-to-layer. For this work, the camera was windowed to an 80 by 80 pixel array (~130 μm/pixel) to enable high frame rate recording (~2500 Hz) during SLM of the samples. Figure 2 is an illustration of the camera observing the SLM process where it collects time-series radiance data for each pixel in the image array. The measured time-series data from two locations of a layer in the various parameter set sample is plotted in Fig. 2 (b).

The measured time-series data from two locations of a layer in the various parameter set sample is plotted in Fig. 2 (b). The time series data is from the middle of the part (location 1) and the border scan region (location 2). In Fig. 2 (b), differences in the shape, number, and frequency of heating and cooling curves are observable when comparing the data from the two locations. These differences in the time-series data potentially correlate to differences in the local part properties across the part. In this work, thermal features have been extracted from the raw time-series radiance data to compress the data and allow informative correlations to be made with local part properties. Thermal feature extraction compresses the data for a layer from ~100 MB to ~0.003 MB by converting the time- series radiance data, L(x,y,t), to a single feature image per layer, f(x,y).

Thermal feature extraction compresses the data for a layer from ~100 MB to ~0.003 MB by converting the time- series radiance data, L(x,y,t), to a single feature image per layer, f(x,y). This enables efficient processing and storage of SWIR imaging data measured for SLM parts.A voxel based method was used to extract the thermal features from the time-series radiance data. The thermal features extracted in this work were the time above threshold, (cid:306), and the maximum radiance Lmax. The time above threshold and maximum radiance are illustrated on a section of the radiance curve from location 2 in Fig. 2 (c). The time above threshold is the total time a pixel is above the apparent phase transition region. As discussed, extraction of the thermal features from each pixel in the camera array results in a 2D map for each layer of a part. The thermal feature map (time above threshold) for the layer from the various parameters sample is illustrated in Fig. 2 (d).

2. (a) SWIR camera observation of SLM process, (b) time series data from location in center of part cross-section (location 1) and border scan region (location 2), (c) definition of thermal features in time series data with melt pool image, and (d) thermal feature maps showing generation of volumetric data and slicing.2.3. Local Part Property MeasurementsThe local part properties measured in this work included the light-to-dark microstructural feature ratio, LDFR, the microhardness, HV, and the volumetric porosity, (cid:307). Fig. 3 (a) demonstrates the light and dark microstructural features in the etch response of SLM 304L stainless steel (polished to 0.05 μm, electrolytically etched with 40/60 nitric acid). The light and dark features are due to differences in the solidification mode, which is described in [22]. The light feature corresponds to material that solidified as ferrite and then transformed into austenite upon further cooling.

The light feature corresponds to material that solidified as ferrite and then transformed into austenite upon further cooling. The dark feature corresponds to material that solidified as austenite and remained austenite. Mapping the ratio of these features within the part cross-sections provided insight into the solidification mode that was most prominent for a particular region. This was completed through implementation of an image processing framework developed in MATLAB. The Vicker’s microhardness (representative indent in Fig. 3 (b), 4.9 N and 10 seconds) was mapped in the part cross-sections following ASTM E384-17 [23]. The porosity in the volume of samples was determined through processing micro-computed tomography (micro-CT) data. This was accomplished by assuming voxels with intensity values three standard deviations less than the mean intensity value of a solid volume corresponded to porosity.

This was accomplished by assuming voxels with intensity values three standard deviations less than the mean intensity value of a solid volume corresponded to porosity. The micro-CT data for the parts was down-sampled to the resolution of the SWIR imaging data by determining the percentage of1615porosity voxels in a 130 μm by 130 μm by 50 μm volume. This process is illustrated for a slice of the various parameter set sample micro-CT data in Fig. 3 (c).(a)d 2(b)Light FeaturesDark Featuresd 1(cid:24)(cid:19)(cid:3)(cid:541)(cid:80)(cid:21)(cid:24)(cid:3)(cid:541)(cid:80)1 mm010 Porosity [%]Fig. 3. (a) Optical micrograph of 304L stainless steel showing light and dark microstructural features, (b) micrograph of indent for determining microhardness, and (c) micro-CT slice with processing to determine voxel based local porosity.3.0.

Local Part Properties CorrelationsThe following details the development of the correlations between local part properties and thermal feature measurements using data from the various parameter set sample. The LDFR and microhardness were determined for the cross-section of the part, while the porosity was measured for the entire volume. Figure 4 contains the data maps for the cross-section of the sample. Figure 4 (a) is the etched micrograph of the sample that shows differences in the amount of porosity and etch response for the different fluence inputs (values annotated to the left). The LDFR plotted in Fig. 4 (b) shows that higher fluence values corresponds to more of the light feature within a section. The increase in the light feature is due in part to differences in thermal boundary conditions during and after solidification. The increase in light feature is also because more energy input leads to larger melt pools and more overlap of laser scans.

The increase in light feature is also because more energy input leads to larger melt pools and more overlap of laser scans. This leads to more re-melting within the parts which corresponds to an increase in the light feature. The same effect is observed for the border scan regions of parts. The microhardness for the part mapped in Fig. 4 (c) appears qualitatively uniform except in areas corresponding to low fluence values where the part was softer, or a pore led to no measurement. The porosity from micro-CT scanning of the part is displayed in Fig. 4 (d) and shows that low fluence values resulted in an increase in porosity. The time above threshold and maximum radiance thermal feature reconstructions are included as Fig. 4 (e) and Fig. 4 (f), respectively. Qualitative comparisons of the thermal features and local part properties shows that low feature values correspond to low LDFRs, low microhardness, and high porosity.

r...-. ··r. ·✓-- · ·Fig. 4. (a) Optical micrograph after etching, (b) LDFR, (c) microhardness, (d) porosity, (e) time above threshold, and (f) maximum radiance in cross-section of various process parameters sample.Figure 4 shows qualitatively clear relationships between local part properties and thermal features. The local part property values and thermal features have been averaged for the various sections of the sample for correlation development through data-driven modeling. The correlations between the LDFR and thermal features are plotted in Fig. 5. Each data point for the LDFR is the average and standard deviation from ~600 pixels. The corresponding thermal feature values are the averages and standard deviations from ~7000 voxels. The LDFR as a function of thermal features was fit to a second order polynomial.

5. Correlation of average LDFR with (a) average time above threshold and (b) average maximum radiance.Figure 6 is the correlation of the averaged microhardness from the various energy inputs with the corresponding thermal feature measurements. Each microhardness value is the average and standard deviation from 20 measurements. Thermal feature values are again from ~7000 voxels. In this figure, the saturation of microhardness is apparent with the additional detail of a slightly decreased microhardness for some of the higher thermal feature values. The relationship between microhardness and the thermal features were fit to a lumped capacitance model.

7. Correlation of porosity determined from micro-CT data with (a) average time above threshold and (b) average maximum radiance.4.0. Prediction of Local Part PropertiesModeling the relationships between thermal features and the local part properties permits predictions for new parts through SWIR imaging measurements. This is demonstrated in Fig. 8 where the models developed in the previous section are applied to the time above threshold thermal feature measurement for the embedded text sample. In this figure, local part property predictions are made for the entire volume of the part. The measured thermal feature predicts significant variation in the local part properties for the sample in the locations corresponding to the embedded text.

- -I"' .., ~010 Porosity [%]20150 50 Microhardness Fig. 8. Demonstration of local part properties predictions using models developed in correlations with thermal feature measurements.50 0250Figure 9 contains the time above threshold prediction of the local part properties plotted next to the experimental measurements. These results are for the cross-section of the embedded feature sample (micrograph included for reference). Figure 9 shows that the predicted areas of1619increased LDFR corresponding to the text border areas match the results found in experimental measurements. The time above threshold was also able to predict the decrease in microhardness for areas corresponding to the text interior. The porosity predictions from the time above threshold show localized areas of high amounts of porosity for the text interiors. There is also the prediction of a slight increase in porosity at the location of the text borders.

In this figure, the LDFR and porosity predictions have increased noise. There are also a larger number of predicted locations around the text that are softer in the maximum radiance predictions.16201 mmMicrographLight-to-Dark Feature Ratio 0 000 1 2 222 3 333 4 444 5Microhardness 150 250 500Porosity [%]1020,. -, . ... : ,_.... . . : • .Lmax PredictionExp.Lmax PredictionExp.Lmax PredictionμCTFig. 10. (a) Etched micrograph of embedded feature specimen cross-section and comparison of maximum radiance predictions with experimental maps of (b) LDFR, (c) microhardness, and (d) porosity.The error of this method has been quantified by subtracting the experimental results from the thermal feature predictions. The error maps for both the time above threshold and maximum radiance predictions are plotted with the experimental local part property maps in Fig. 11. These maps shows that the thermal features predict the properties of the part with some error.

11. These maps shows that the thermal features predict the properties of the part with some error. There is an increase in error around the embedded text, but the increase is tolerable. To demonstrate this, histograms of the error were generated for the cross-sections of LDFR and microhardness predictions and for the volumes of the porosity predictions.1621Light-to-Dark Feature Ratio 0 000 1 111 2222 3 333 4 444 5(a)LDFR Error 0 -5 5Microhardness 500 150 0 250ness Microhardness Error -50 050(b)Porosity [%] 0 10 20 0(c)Porosity Error [%] -20 0 202 mm... - 4---w..:: •.,-..:: ·-· ~:· ~ ...::: ~...Lmax Error Exp. (cid:306) Error Error = Prediction - Experimental Fig. 11.

The error for microhardness is offset from 0 due to the models developed in Fig. 6. In the models from Fig. 6, the microhardness saturates at a value of ~222 and this is ~7 hardness values less than the mean for the surrounding of the embedded feature sample. Figure 12 (c) shows that a majority of porosity error was between 0 and 2% for both thermal feature predictions (error over ~580,000 voxels). The time above threshold also performed better for this local part property.

The time above threshold also performed better for this local part property. It should be noted that a reduction in the reported error is possible by improving the indexing of predictions and experimental methods and with the development of stronger models for the relationships between thermal features and local part properties.16226015]%[s e x Pli504030(a) (cid:306) Lmax]%[s e x Pli10(cid:306) Lmax(b)t n e c r e P20t n e c r e P5100 -54321 1 0 LDFR Error23450 -50 -40 -30 -20 -100 Microhardness Error10 20 30 40 5080]%[60(c) (cid:306) Lmaxs e x o Vl40tn e c r e P200 -50 Porosity Error [%] Fig. 12. Error distribution for thermal feature predictions of (a) LDFR, (b) microhardness, and (c) porosity.432112345In Figs.

12. Error distribution for thermal feature predictions of (a) LDFR, (b) microhardness, and (c) porosity.432112345In Figs. 9 and 10, both the time above threshold and maximum radiance were able to qualitatively predict the variances in the local part properties in the embedded feature sample. Quantifying the error of the predictions in Figs. 11 and 12 showed that the time above threshold performed the best for the LDFR and porosity. This is because the time above threshold accounts for the total amount of time the radiance of a voxel is above the apparent phase transition region. This is a strong thermal feature to use in predicting the LDFR since the microstructural feature is related to the total number of times the material is re-melted. The time above threshold feature is also better for porosity predictions because the more time a voxel is above the apparent phase transition region the more likely it is to correspond to fully dense material.

Additionally, the increase in porosity due to keyholing was related more strongly to the time above threshold which led to better predictions. In general, the increased error of maximum radiance predictions is due to weaker correlations with the local part properties and a larger amount of noise in the measurement. The errors of the microhardness predictions were within the standard deviations of the experimental measurements corresponding to the extreme energy inputs in Fig. 6 for both thermal features. The results in Figs. 8-12 show that predictions of local SLM part properties can be made with tolerable error for complex features in SLM parts.5.0. Summary and ConclusionsIn this paper, a voxel based framework was used to process SWIR imaging data measured in-situ and make predictions of SLM 304L stainless steel part properties on a local basis.

Thermal features including the time above threshold and maximum radiance were extracted from time- series SWIR imaging data and correlated with the local part properties. The time above threshold1623was found to correlate the strongest when developing models of the local part properties as a function of the thermal features. The developed models were used to make predictions of the local part properties based on new thermal feature measurements for an SLM sample manufactured with an embedded feature and the error was quantified. The time above threshold outperformed the maximum radiance in making predictions of the LDFR and porosity, but performed similarly for microhardness.The errors of the local part property predictions could be reduced by moving towards machine learning algorithms with the incorporation of multiple thermal features that capture different dynamics of the SLM process.

The framework reported in this paper can be used to qualify parts, or flag them for further ex-situ analysis, on a local basis with further development. In this further development, confidence intervals for the minimum size of defective local part property variation predictable will be defined.AcknowledgmentThis work was funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.References[1] M. Grasso and B.M. Colosimo, Process defects and in situ monitoring methods in metal powder bed fusion: a review, Measurement Science and Technology 28 (2017) 1-25. [2] Y. Zhang, J.Y.H.

viewed PaperIn-situ Modification of a High Entropy Alloy With 2.4% Molybdenum Using LPBF, and its Effect on Microstructure and Corrosion ResistanceM. Ritchie*, S. Mehraban, S. G. R. Brown, D. Butcher, J. Cullen, M. Calvo-Dahlborg and N. P. LaveryAbstractLaser powder bed fusion (LPBF) components display higher porosity compared to parts made by conventional processes and these pores act as preferential initiation sites for pitting corrosion to occur. In stainless steels such as 316L, molybdenum is 3.5X more effective at enhancing the pitting resistance than chromium, without adding unwanted nitrides to the alloy. In this work, the effect on corrosion resistance is reported for an Al-Cr-Mn-Ni-Fe high entropy alloy (HEA) gas atomised specifically for LPBF, as well as the effects of modifying the alloy by blending the HEA with molybdenum.

In-situ LPBF processing, even for low levels of additions has made the comparison difficult, as the pitting resistance is so strongly linked to the porosity, which is higher in the in-situ process. Pitting resistance for both the original HEA and the doped HEA will be compared between samples processed by casting and by LPBF.Keywords: High Entropy Alloy, Laser Powder Bed Fusion, Pitting Corrosion.IntroductionThe LPBF process continues to grow as an important additive manufacturing (AM) process for a diverse set of critical applications where high density is required. Porosity occurring from the LPBF process has been shown to reduce the mechanical properties of metals, from tensile strength and elastic modulus to elongation [1–4]. This is particularly relevant in aerospace and automotive industries where failure in critical components could result in loss of life.

This is particularly relevant in aerospace and automotive industries where failure in critical components could result in loss of life. Aerospace manufacturing is therefore subject to extensive quality control, a hurdle for the AM industry as many studies report disparities in mechanical properties due to the many parametrical influences on component quality [3,5–7]. Unlike other ways of manufacturing net-shape parts, LPBF offers much greater design freedom, with an increasing palette of alloys, mostly based on traditional alloys used in specific applications [8,9]. These alloys were developed a long time ago for other processing routes such as casting such as AlSi10 and AlSi7 or forging with Ti-6Al-4V, where the cooling rates are much lower than LPBF.

These alloys were developed a long time ago for other processing routes such as casting such as AlSi10 and AlSi7 or forging with Ti-6Al-4V, where the cooling rates are much lower than LPBF. Even when traditional alloys can be used, the high cooling rates result in heterogeneous microstructures with long columnar grains, leading to different mechanical properties to what might be expected, such as higher hardness at the expense of lower ductility [10,11]. An added consideration is the porosity levels, which although increasingly low in new generations of machines, still affects properties such as corrosion and fatigue strength, with surface pores operating as preferential sites for pitting corrosion [12,13].

However, not all alloys can be successfully used in LPBF, and those with poor weldability such as Ni based alloys like Hastelloy X often having the narrowest process operating windows making it difficult to make parts without defects, namely porosity and/or solidification cracking [14].H13 is a good example of an alloy which is difficult to process by LPBF. As a steel used in tooling such as injection moulding it is expected to have good thermal fatigue strength and an acceptable level of corrosion resistance. H13 is currently being modified to process better, both during LPBF and after building, with integrated heat treatments [15–17]. Because of this, there is an increasing body of work going into deriving AM specific alloys (e.g. Scalmalloy aluminium alloy), or modified traditional alloys (e.g. H13) [18,19].

Because of this, there is an increasing body of work going into deriving AM specific alloys (e.g. Scalmalloy aluminium alloy), or modified traditional alloys (e.g. H13) [18,19]. However, there373are still many areas where more work needs to be done such as alternatives to high grade aluminium alloys.HEAs are generally accepted to be a system of 5 or more principal elements, the concentrations of which fall between 5 and 35at.% [20–22]. This differs from the standard makeup of stainless steels for example, which use Fe to make up the bulk of the composition, with additional elements added in smaller quantities to provide preferential characteristics. The compositions of HEAs can be equimolar or not, either way, the vast number of possible combinations is one of the main reasons for the interest in the field along with the simplicity of their design and the ability to tailor constituent elements to achieve desired characteristics [20].

The majority of HEA creation has occurred in arc melters or induction coils, and LPBF is only just being researched as an alternative manufacturing method. HEA production can be a complex process with high levels of control required to manufacture them. Therefore, the parameter control and fast cooling rates generated by LPBF suggests that the process could be suitable replacement to create components from HEAs [20]. Work on HEAs processed by laser cladding and LPBF is increasingly being published, but this method of production is still in its infancy [20,23].The pitting resistance equivalent number (PREN) is used to predict a stainless steels resistance to pitting corrosion, with it equating to the wt.%Cr + 3.3 × wt.%Mo + 16 × wt.%N.

The addition of 2.4% Mo to the HEA in this paper was based on its theoretically larger influence on the pitting resistance than Cr [24,25].Materials and MethodologyThe 316L and HEA1 analysis and corrosion specimens were produced from virgin 15µm - 45µm powder provided by Sandvik in a Renishaw AM400 machine on full sized build plates after running parameter optimisations studies. To create the HEA1 + 2.4% Mo mixture, 122.951g of powdered Mo (<5µm) was added to 5Kg of HEA1, and put into a LFA V-type blender for 2 hours at an agitation speed of 60r/min to allow for homogenous distribution of the Mo. Compositional analysis was undertaken on the resultant blend to confirm the target levels of Mo had been reached before it was loaded into a reduced build volume (RBV) within the AM 400 to produce samples for analysis and corrosion using parameters based on those used for HEA1.

Table 1 shows the target and achieved compositions of the blended mixture with Figure 1 showing SEM images of the powder at 550x (a) and 1100x (b) magnification. HEA1 particles appear much larger due to the size of the particles used.Table 1. Target and achieved compositions for HEA1 and HEA1 + 2.4% Mo (all in wt.%)Fe 33.97Cr 27.3Ni Mn 10.3 25Al 2.8N 0.20Si 0.4C Mo0.02 Target HEA1+2.4 Mo 33.15 26.64 24.40 10.05 2.73 0.20 0.39 0.02 Achieved HEA1+2.4 MoHEA1 (Measured)2.431.45 26.39 25.36 10.46 3.20 0.24 0.54 0.03 2.33374MoHEA1(a)(b)Figure 1. HEA1 + 2.4% Mo at 550x (a) and 1100x (b) magnificationCast samples were created from a combination of feedstock and weighed approximately 40g. These feedstocks were re-melt samples produced through LPBF, or powder weighed out and compacted using a Baileigh hydraulic press. All the feedstock produced stemmed from the same 15µm - 45µm powder referenced previously, as did the Mo.

All the feedstock produced stemmed from the same 15µm - 45µm powder referenced previously, as did the Mo. The feedstock was melted under argon (>99.9% pure) in an induction heater in 15mm diameter crucibles. Once molten, the 316L was held at temperature for 1 minute whilst the HEA samples held for 5. The extra duration for the HEAs was to allow the elements to diffuse uniformly through the sample. Two samples of each alloy were produced, one to section for compositional, microstructural, and porosity analysis, and another for corrosion testing.All samples for analysis were subjected to a 5-stage grinding and polishing regime, resulting in a 1µm finish. Porosity analysis was conducted on AM samples using a Biolin Scientific Sigma 700 force tensiometer and Archimedes calculations. All samples were then imaged using a Zeiss Observer inverted light microscope, with its stitching function used to map the full sample area.

All samples were then imaged using a Zeiss Observer inverted light microscope, with its stitching function used to map the full sample area. The images were put through ImageJ to analyse the porosity and in the case of the AM samples, compare this data to that obtained from the tensiometer. Microstructure and compositional analysis were performed on a JEOL JSM-6010 Secondary Electron Microscope with Energy Dispersive X-ray Spectrometry SEM-EDS. EDS maps were used to verify the bulk chemistry with the target chemistries with point analysis used to acquire the compositions of different phases within the microstructure. All SEM and EDS images used are taken of horizontal slices, in the case of the AM samples, these were perpendicular to the build direction.Corrosion specimens of approximately 1.5mm thickness were cut from the AM and cast samples, soldered to electrical wiring and subsequently cold mounted. These were then polished as previously discussed.

These were then polished as previously discussed. Once complete, they were exposed to an open circuit potential (OCP) scan and a cyclic polarisation test, with multiple tests were carried out per sample.ResultsMicrostructure AnalysisFigure 2 shows the cast microstructures of HEA1 (a) and HEA1 + 2.4% Mo (b). These images show that initially HEA1 has two dominant phases, and the addition of the Mo creating375a third. Thermocalc predicts that the lighter phase in HEA1 is FFC_L12, with the darker region being BCC_B2.(a)(b)Figure 2. Cast microstructure of HEA1 (a) and HEA1 + 2.4% Mo (b)The EDS mapping of HEA1 in Figure 3 shows a high Cr level present in the darker BCC phase, which coincides with a lower Ni reading. The inverse is true in the lighter FCC phase, with the Ni being comparatively higher than the BCC phase and Cr levels dropping.

The inverse is true in the lighter FCC phase, with the Ni being comparatively higher than the BCC phase and Cr levels dropping. Point scans of each phase confirm this observation, with the Cr and Ni rich phase readings being 12.98wt.% and 13.04wt.% higher respectively than their depleted counterparts. These readings are similar to those predicted by Thermocalc.(a)(b)Figure 3. Cr (a) and Ni (b) EDS maps for cast HEA1As with HEA1, the EDS analysis of HEA1 + 2.4% Mo in Figure 4 shows Cr abundance and Ni depletion and in the BCC phase, with the inverse true in the FCC phase. The addition of Mo does not appear to have affected this relationship, with the EDS map of Mo suggesting a similar phase relationship to Cr, with a greater abundance in the BCC phase compared to the FCC. The third phase not apparent in HEA1 appears to be another that is richer in Cr and Mo and depleted of Ni.

The third phase not apparent in HEA1 appears to be another that is richer in Cr and Mo and depleted of Ni. Again, point scans of these areas confirm this prediction, with an even larger compositional differential between the phases when compared to HEA1 observed. Analysis shows a 15.94wt.% and 2.15wt.% increase in Cr and Mo levels respectively in the BCC phase compared to the FCC. Conversely, Ni is observed at 15.81wt.% higher in the FCC phase compared to the BCC. The elements in the third phase occur at levels similar to those in the BCC with the other elements remaining relatively constant between each phase.376(a)(b)(c)(d)Figure 4. Cast microstructure (a) and EDS maps of Ni (b), Cr (c), and Mo (d) for HEA1 + 2.4% MoFigure 5 shows the AM microstructures of HEA1 (a) and HEA1 + 2.4% Mo (b).

Cast microstructure (a) and EDS maps of Ni (b), Cr (c), and Mo (d) for HEA1 + 2.4% MoFigure 5 shows the AM microstructures of HEA1 (a) and HEA1 + 2.4% Mo (b). Comparing these to the microstructures seen from the cast specimens of the same alloys in Figure 2, the dendritic structure looks to have disappeared at 500x magnification, leaving behind more defined grain boundaries. The white spots seen in (b) are likely due to charging from the electron beam.(a)(b)Figure 5. Microstructure of AM HEA1 (a) and HEA1 + 2.4% Mo (b)377Similar to the cast microstructures, different phases appear to occur in the microstructure, however, EDS mapping and point scans show that these different regions are compositionally comparable, with a maximum elemental deviation of 0.28wt.% between each region. EDS maps of Cr and Ni are where this is seen most prominently as shown in Figure 6.

EDS maps of Cr and Ni are where this is seen most prominently as shown in Figure 6. When compared to their cast counterparts in Figures 3 and 4, there are no distinct areas of richness, depletion, or segregation as there were in the cast samples.(a)(b)(c)(d)Figure 6. EDS maps of Cr and Ni for AM HEA1 (a,b) and HEA1 + 2.4% Mo (c,d)Corrosion AnalysisFigure 7 shows the result of cyclic polarisation scans on cast and AM samples of 316L, HEA1 and HEA1 + 2.4% Mo. First observations show that, whilst the 316L test specimens have a clear pitting potential (Epit), the HEA alloys corrode at a steadier rate with a much less obvious Epit. Other graphed aspects between 316L and the HEAs can also be compared, such as the metastable pitting phase that occurs after the corrosion potential (Ecorr) is reached. With 316L, the current measured remains fairly consistent as the voltage applied continues to increase, until the Epit is reached.

With 316L, the current measured remains fairly consistent as the voltage applied continues to increase, until the Epit is reached. The HEAs current however consistently increases, with only minor evidence of any metastable pitting. Key data from these graphs has been extracted and is displayed in Table 2.378(a)(b)(c)Figure 7. Cyclic polarisation graphs for cast and AM samples of 316L (a), HEA1 (b), andHEA1 + 2.4% Mo (c)Table 2.

Cyclic polarisation graphs for cast and AM samples of 316L (a), HEA1 (b), andHEA1 + 2.4% Mo (c)Table 2. Corrosion data extracted from Figure 7316L Cast 316L AM HEA1 Cast HEA1 AM HEA1 + 2.4% Mo Cast HEA1 + 2.4% Mo AMEcorr (V)0.058 -0.047 -0.199 -0.254 -0.2740.208Epit (V)0.376 0.927 0.231 -0.047 0.1000.159Epit - Ecorr (V) 0.434 0.974 0.430 0.207 0.3740.367Re-passivation potential (V) -0.191 -0.183 -0.208 -0.294 -0.2980.324DiscussionThe differences in microstructure observed between the AM and cast samples imaged, as well as the differences in phase formation and elemental segregation are often attributed to the rapid rate at which the AM process heats and cools the part. The HEA castings were held at molten for 5 minutes to allow for full diffusion of the elements within the alloy, and the cooling happens over a number of minutes.

The HEA castings were held at molten for 5 minutes to allow for full diffusion of the elements within the alloy, and the cooling happens over a number of minutes. The AM process conducts the fusion of an entire layer of a density cube within a second, with the heat created being rapidly dissipated through the adjacent, unmelted powder and the baseplate. This facilitates minimal diffusion and subsequently, no areas of rich and depleted elemental abundance like that which is seen in the cast samples.The addition of 2.4wt.% Mo to HEA1 does appear to have added a third clear phase to the microstructure, although further work with Thermocalc and X-ray diffraction (XRD) is379needed to confirm the details of this phase, and why the Mo appears to migrate towards the Cr rich regions. Along with this third phase, the phases in HEA1 + 2.4% Mo appear to be more irregular than those in HEA1.

Along with this third phase, the phases in HEA1 + 2.4% Mo appear to be more irregular than those in HEA1. The phases of HEA1 are observed to form parallel or perpendicular to each other, whereas there is more of a branch-like structure with the Mo addition. This, however, could be due to processing, as every cast is different, so further repeats and analysis would be required to verify whether it is infact a characteristic induced by the addition, or simply variations in the manufacturing process.HEA1 + 2.4% Mo built in the RBV was more difficult to process than anticipated, but this could be down to multiple factors. Mo has a melting point of over 2600°C, compared to the alloyed HEA1 which is below 1800°C. The Mo powder was also not manufactured specifically for AM, and with its particle size being <5µm and with that of HEA1 being between 15µm - 45µm, the Mo could have increased the packing density, and therefore the amount of energy required per mm3.

These combined suggest increasing the energy density when processing the alloy through AM could overcome the melting point differential, and any issues caused by the packing density. Also, due to the sphericity of the particles not being a target for the Mo in production, the subsequent lack of could decrease the flowability of the overall material when distributed by the wiper blade, leading to problems caused by an uneven supply.The corrosion of the HEAs appear to translate well between the different manufacturing processes when compared to that of 316L, which has a larger differential. In the regions between Ecorr and Epit for the HEAs and 316L, clear metastable pitting occurs in 316L along with minimal change in current, but with both HEAs, the current continues to increase in line with the voltage, with only marginal signs of the metastable phase.

The metastable phase occurs due to the formation of a passive film on the surface of the metal, protecting the corrosion site from further attack. All the alloys have Cr present in similar quantities which is one of the elements responsible for the generation of the passive layer, however it appears more effective during that stage of the corrosion cycle in 316L than the HEAs. Having said this, the Epit - Ecorr for the three cast samples shown in Table 2 do not differ significantly, ranging from 0.434V for 316L to 0.374V for HEA1 + 2.4% Mo.The Mo addition to HEA1 does not appear to have had the desired effect when comparing the cast specimens, with the Epit – Ecorr actually decreasing by 0.056V. This could be due to the greater disparity between the Cr rich BCC phase and its depleted FCC phase, increasing from 12.98% in HEA1 to 15.94% in HEA1 + 2.4% Mo.

This could be due to the greater disparity between the Cr rich BCC phase and its depleted FCC phase, increasing from 12.98% in HEA1 to 15.94% in HEA1 + 2.4% Mo. With the Mo segregation also following this trend there are regions of the microstructure that have a significantly lower abundance of the elements associated with corrosion resistance. These depleted regions could therefore be acting as a preferential site for corrosion to occur. Conversely, when the addition is processed through the AM route, it does seem to have a preferential effect on the corrosion resistance, with the Epit – Ecorr rising by 0.160V. The EDS analysis of the AM specimens showed a much more homogenous elemental distribution, and therefore no regions that would act as a preferential site of corrosion when compared to the cast samples in the same fashion.

This could explain why the Mo addition benefits the corrosion resistance in the AM sample, whilst not being seen in that of the cast.ConclusionTo conclude, the microstructural changes by adding Mo to HEA1, including the third phase formation and the migration of Mo to the Cr rich phases in the cast specimens needs to380be understood to a greater extent, as well as its effect on the corrosion performance. This is less relevant in the AM specimens where no significant depletion occurs in any of the grains or phases. The corrosion data extracted from the cyclic polarisation scans for the HEAs suggests a varying effect of the Mo addition on their pitting resistance that is dependent on their manufacturing method. However, the irregular nature of cast and AM surfaces necessitates further analysis on HEA1 and HEA1 +2.4% Mo to determine the effectiveness of the Mo addition at increasing the corrosion resistance of HEA1.[1] Kan WH, Chiu LNS, Lim CVS, Zhu Y, Tian Y, Jiang D, et al.

This research aims to develop a method of low-cost and high-resolution thermal monitoring system using near-infrared (NIR) wavelength band emission monitoring to detect anomalies. This research will compare more expensive infrared (IR) wavelength band monitoring methods to the cheaper NIR method and other drawbacks brought about by monitoring one wavelength band over the other.IntroductionThere exist many methods of monitoring L-PBF that can be implemented for several types of equipment in a vast variety of printable materials. One such monitoring method is the collection of specific wavelength emissions during and after the melting and solidification process for L-PBF with atomized metal powder.

One such monitoring method is the collection of specific wavelength emissions during and after the melting and solidification process for L-PBF with atomized metal powder. This can provide information about the health of the equipment, but more importantly, this can provide information that correlates with the quality and properties of the printed material.For this experiment, a hypothesis was formed that by using NIR process monitoring of the build surface during and after the printing of materials, information about the quality and properties of the final material can be correlated with the collected NIR sensor information. To assess this hypothesis, a thermal testbed was initially set up to prove the ability to collect thermally correlated data using a NIR camera.

To assess this hypothesis, a thermal testbed was initially set up to prove the ability to collect thermally correlated data using a NIR camera. The equipment from this testbed setup was then moved into the L-PBF system for testing in simulated and actual printing environments.This hypothesis was based upon the knowledge of NIR signal emissions of black bodies at known temperatures, as well as a NASA patent that was exploring the ability to correlate NIR and IR signals from L- PBF processes to re-construct a 3D representation of the build detect thermal anomalies, in the form of visually identifiable “hot-spots” [1]. The setup for the NASA research can be seen in Figure 1.

The setup for the NASA research can be seen in Figure 1. Initially, the experiment outlined in this paper was aimed toward replicating this setup, however, it was seen that using only the NIR camera was still able to yield meaningful results without the need for an infrared camera.The use of IR cameras in L-PBF is also not uncommon and has been proven as a viable method of process monitoring [2]; however, common IR cameras are very expensive, hard to connect, and tend to have low- resolution sensors and slow frame rate for image capturing. These challenges have all been seen through past research experiences using IR cameras. By using a NIR camera as the main process monitoring sensor, it should allow for a less-expensive, higher-resolution, and easier image capture capabilities.

By using a NIR camera as the main process monitoring sensor, it should allow for a less-expensive, higher-resolution, and easier image capture capabilities. This should allow for more granular process monitoring during the build cycle, including the ability to adjust collection parameters to detect different forms of anomalies, which is not possible using traditional IR cameras.A common IR camera used in our research is the FLIR® A35 which yields a resolution of 320 by 256 pixels, 60 frames-per-second, and a resale cost of approximately $6000 at the time of writing this manuscript. The NIR camera used for this experiment was the Pixelink® PL-D734MU-NIR-T which yields a resolution of 2048 by 2048 pixels, 90 frames-per-second, and a cost of $2200. By using the NIR camera for process monitoring, the smaller nominal pixel size of the sensor allows for geometrical accuracies which are far more1234meaningful than those possible with standard IR cameras.

The drawback of using a NIR camera over an IR camera is the loss of temperature accuracy. This research also aims to prove that, in L-PBF monitoring, an accurate temperature is not as valuable as the detection of temperature variations for correlation with certain process defects, such as bulk porosity; therefore, using a NIR camera is a useful method of process monitoring. [3], [4]Figure 1: Setup image from US Patent 10,688,560 L-PBF IR and NIR process monitoring setup. [1]Experimental SetupThere existed two main experimental setups for this research. The first consisted of the monitoring equipment and a simulated heat source in a controlled lab environment. This was done initially to decide upon the validity of using a NIR camera as a method of thermal process monitoring.

This was done initially to decide upon the validity of using a NIR camera as a method of thermal process monitoring. Once this capability of the NIR camera was confirmed and configured, the same equipment was then installed into the L-PBF system for both simulated and actual printing processes to explore its use as a process monitoring system in the additive manufacturing environment.Testbed ExperimentationThe first testbed setup consisted of the Pixelink NIR camera with a band-pass filter of 700 nm, a FLIR IR camera, a standard ceramic hotplate capable of reaching 500°C, and various test artifacts of dissimilar materials and surface finishes. This setup can be seen in Figure 2. This testbed setup allowed for the selection of a proper camera configuration with exposure time and camera aperture.

This setup can be seen in Figure 2. This testbed setup allowed for the selection of a proper camera configuration with exposure time and camera aperture. The band-pass filter was selected to filter all non-useful wavelengths from the laboratory background while still allowing for the collection of useful wavelengths generated by the heat source and test artifacts.1235Figure 2: Thermal testbed setup for initial NIR and IR camera exploration.The initial testbed setup performed to explore the use of the NIR camera consisted of two main experiments. The first was the use of a test piece of aluminum with a reflective surface. The second was the use of multiple dissimilar materials of varying emissivity. For the first experiment, the aluminum piece was placed on the hotplate and heated from room temperature to 500°C. The NIR and IR cameras were set to record over the duration of the heating.

For the first experiment, the aluminum piece was placed on the hotplate and heated from room temperature to 500°C. The NIR and IR cameras were set to record over the duration of the heating. Upon reaching the set temperature, the IR camera showed an increase in detector intensity on the hotplate and a less-significant increase of intensity seen on the aluminum sample as seen in Figure 3. This can be explained by the differences in emissivity of the hotplate and the aluminum piece.During the second experiment, four different test pieces were run through an identical experiment as in the first experiment. The four pieces consisted of pieces made of matte aluminum, brass, black-oxide alloy steel, and reflective aluminum. Once at the set temperature, the brass, steel, and matte-finished aluminum, as well as the hotplate surface all, showed an increase in detector intensity on the IR camera, while the reflective aluminum showed little intensity change like in the first experiment.

Looking at the NIR camera signals, it can be seen there was an increase in detector intensity around the reflective aluminum piece as well as the hotplate surface surrounding it. These results can all be seen in Figure 3.1236Figure 3: Thermal testbed preliminary results.After the preliminary results of the testbed experiments showed a qualitative correlation between NIR detector intensity and thermal conditions, a third experiment was designed to quantitatively analyze these observations. This last experiment consisted of the same hotplate and the aluminum piece from experiment one and followed the same heating set point. The aluminum piece was not of a known temperature as a thermocouple was not installed on this system. The hotplate temperature is known by using the internal thermocouple of the hotplate, used for heating control, and can be compared to the recorded detector values of both the NIR and the IR cameras.

The hotplate temperature is known by using the internal thermocouple of the hotplate, used for heating control, and can be compared to the recorded detector values of both the NIR and the IR cameras. This was done by comparing the hotplate temperature to a set point on the plate surface, as seen in Figure 4. The detector values from this point on both cameras were recorded over the length of the heating cycle. The results can be seen in Figures 5 and 6.Figure 4: Hotplate set point location for both IR and NIR cameras.1237Figure 5: Comparison of the hotplate temperature to IR detector results.Figure 6: Comparison of the hotplate temperature to NIR detector results.The IR signal data corelates well to the hotplate temperature, while the NIR signal begins to exponentially increase in intensity after about 100 seconds, or around 200-300°C. Towards the end of the heating cycle, the NIR signals appear to oscillate near the 500°C setpoint.

Towards the end of the heating cycle, the NIR signals appear to oscillate near the 500°C setpoint. When comparing back to the NIR recorded frames, the plate stays close to the setpoint temperature, however the heating element is occasionally seen emitting brighter and dimmer, most likely due to the heating control unit of the hotplate regulating the setpoint temperature. Through this experiment it can be seen that a NIR camera can detect thermal changes and can be further explored within the L-PBF system.1238System Installation and ConfigurationBoth the NIR and IR cameras were installed in the Farsoon FS271M L-PBF system where the NIR camera was mounted outside the system with a view through a fused silica window capable of transmission wavelengths in the range from 200 to 2200 nm. The IR camera was installed inside the system with no other filtering.

The IR camera was installed inside the system with no other filtering. Both cameras were viewing the build plate at an angle of between 15-30 degrees, and at a distance like the testbed setup which was around 50cm. The setup can be seen in Figure 7.Figure 7: System installation photos.Due to the high-intensity emissions generated by the laser, a new filter is needed to remove excess emissions to yield a more meaningful image of the printing process. To yield the most meaningful data, without saturating the NIR sensor, an optimal combination of exposure time and filter configuration was needed. To find the most optimal range, 30 different filter and exposure time configurations were tested, all of which are outlined in Figure 8.In each of the experiments detailed below, the laser was fired on a bare build plate with nominal build parameters.

There was a series of 8 squares positioned on the build plate to allow for image correction and to identify any possible locations for laser reflection. The filter configuration of the camera as well as the exposure time was changed for each of the experiments as a single layer was deposited without powder on the stationary build plate. The NIR camera was set to record the entire duration of laser-firing. These data were exported from the system as a series of image frames.Using Python and OpenCV, the sequence of images was analyzed and the intensity of each pixel location in the subsequent image was summed into a matrix. This matrix is the total intensity experienced1239by a select NIR pixel throughout the entire laser-firing.

This matrix is the total intensity experienced1239by a select NIR pixel throughout the entire laser-firing. These results are shown as the color images in Figure 8.Figure 8: System simulated experiment; varying exposure times and NIR camera filters.An exposure time between 30 and 100 microseconds with a filter of ranges 813 – 948nm, 747 – 757nm, or 900 – 950nm (bottom 3 filters and leftmost 3 exposure times) yield an image without excess reflections and ghosting, as seen in Figure 9 on the left- and right-most images. These reflections can be seen much more dramatically in higher exposure times and less restrictive filters, which overwhelms the laser intensity data intended to be captured. The configuration selected from this set of experiments to continue with for further experimentation was an exposure time of 60 microseconds and a filter of range 900 to 950 nanometers.

The configuration selected from this set of experiments to continue with for further experimentation was an exposure time of 60 microseconds and a filter of range 900 to 950 nanometers. This configuration was selected as it yielded a high-rate capable exposure time and did not over-saturate the sensor with ghosting or reflections.1240Figure 9: Over-exposed (left and right) and ideal (center) filter and exposure time configurations.Figure 10: Single frame NIR captured data from selected configuration (center) and over-exposed configurations (left and right).Another aspect of this research was to figure out if the NIR camera can detect inter-layer temperature, or the temperature of the powder bed after the completion of a single layer. In order to detect inter-layer temperature, the NIR sensor must be capable of exposing for extended amounts of time.

In order to detect inter-layer temperature, the NIR sensor must be capable of exposing for extended amounts of time. It was seen that an exposure time beyond 5 seconds with no background NIR interference, taken within 2 seconds of the laser firing on the build plate showed no detectable increase in intensity. Due to the time between laser-firing between layers being 13 seconds, with the powder wiper bar of the Farsoon obstructing the build plate view for most of this time, it was decided that using the NIR camera for post-solidification thermal monitoring was not plausible using the current setup. Future experiments focused on the analysis of the laser spot size data from the frames of the NIR data, as seen in Figure 10.Induced Failure Experiment DesignA test build was developed to induce defects in printed parts by varying the fill laser power part parameter while using all other nominal settings as defined in Table 1.

These powers ranged from 100 to 400 watts in increments of around 75 watts. Parts at 100 watts are designed to yield lack-of-fusion defects while the 400-watt parts are designed to melt and solidify near the peak energy capability of the system.Table 1: Process parameters used for all samples.Parameter Scan velocity Layer thickness Base plate temperature Hatch spacingValue 1000 mm/s 30 µm 100 °C 120 µm1241There are seven unique geometries printed in this build. There were three sizes of 10mm tall cylinders including 5-, 10-, and 15-mm diameter variations as seen in Figure 11, images A, B, and C. There were three parts printed to induce part over-heating as seen in images D, E, and D. Finally, there was a 10mm cube.

There were three parts printed to induce part over-heating as seen in images D, E, and D. Finally, there was a 10mm cube. These geometries were placed evenly spaced on the build plate in as seen in Figure 11, and their respective build parameters in Table 2.Figure 11: System experimental setup.Table 2: Modified fill laser powerRow Number 1 2 3 4 5Modified Fill Laser Power (W) 100 175 225 325 400Experimental ResultsThe induced defect experiment designed above was ran in the Farsoon L-PBF system with the same NIR and IR camera setup decided upon after the lens filter and exposure time selection experiment. The system used nitrogen as the inert gas and powder atomized 316L stainless steel powder as the metal. After the experiment was run, the IR and NIR data were processed and analyzed. These results can be seen in Figure 12, with more detailed processes found in the following sections.

After the experiment was run, the IR and NIR data were processed and analyzed. These results can be seen in Figure 12, with more detailed processes found in the following sections. For scale, the parts in columns labeled d, e, f, g, and b (from Figure 11) are all 10 mm in size and all parts are 10 mm tall.1242Figure 12: System actual printing experiment; raw results for IR and NIR cameras.IR Reference DataThe FLIR A35 was used to collect infrared thermal data to allow for correlation with the NIR collected data for locations of high and low thermal radiation or to further investigate specific changes noted in the NIR data. In Figure 12 you can see a frame captured from an arbitrary layer by the IR camera. As parts are solidified, they appear brighter in intensity and cool over time to a lower intensity.

In Figure 12 you can see a frame captured from an arbitrary layer by the IR camera. As parts are solidified, they appear brighter in intensity and cool over time to a lower intensity. One of the parts on the lower left appears to be higher in intensity than all others near it, this is because it was fired upon more recently in reference to when this frame was captured. IR data is still useful in the process monitoring of L-PBF, and in combination with the NIR data collected from this experiment, more insights into the process may be discovered. This exploration will take place in future research as the NIR and IR data will be combined, and thermal gradients extracted.NIR DataThe NIR camera was set to collect data throughout the entire build with the configuration outlined in the experimental setup. This data was collected as a series of images for each frame recorded by the camera. This image stack was then processed using Python and OpenCV to isolate the specific attributes.

This data was collected as a series of images for each frame recorded by the camera. This image stack was then processed using Python and OpenCV to isolate the specific attributes. It was first filtered to remove pixel values of 5 or lower. It was seen that including these pixel values in a layer-wise sum of the frames contribute to the noise of the camera sensor. These pixel values were removed by overwriting them with a value of 0.The image processing algorithm was used to calculate the total pixel intensity over each layer. For example, if a single pixel was full brightness (255) for two frames in the video, and off (0) for the rest of the video, the final pixel intensity would equal 510. The resulting data can be seen in Figure 12 as the greyscale image on the right. Parts shown on the top in all images of Figure 12 are printed at the highest laser power while parts on the bottom are printed with the lowest. The parts in between are scaled as seen from Table 2.

Parts shown on the top in all images of Figure 12 are printed at the highest laser power while parts on the bottom are printed with the lowest. The parts in between are scaled as seen from Table 2. As laser power increases, the shape of the part becomes clearer, and the amount of spatter also increases. The lower-power parts are more globular in shape and lack spatter; higher-power parts are more defined and present significantly more spatter.To detect changes in thermal signatures better, the image from the NIR cumulative data in Figure 12 was converted to a color map with more dramatic changes visible in higher-value pixels. These results can be seen in Figure 13 where a laser power increase is now shown from left to right.1243Low Laser PowerHigh Laser PowerFigure 13: Analyzed NIR signal information.The 10 mm diameter cylinders were the focus for the rest of the experiment, as they were the simplest geometry that could be reasonably scanned in post-processing equipment.

Future research will be conducted to evaluate the other geometries, specifically the defect inducing geometry.The colormap image and zoomed-in detailed portion seen in Figure 13 clearly show that, in general, the NIR signal data increases in intensity as laser power increases. This can be seen as a change in the colormap scale (Python Jet) where blue is the lowest, and red is the highest intensity. By combining this layer-wise NIR data in series with all other printed layers, it becomes possible to re-create a model of the NIR signals in a spatial reference frame and allow for the ability to identify areas of concern.

While this research timeline was short, the re-creation of each layer’s NIR was not able to be processed, the single layer was used in correlation with other post-processing data collection to find correlations between NIR layer- wise data and part density.CT AnalysisUpon completion of the build, there were 10 samples of the 10 mm cylinders which were further analyzed using x-ray CT. There were two samples for each laser parameter configuration. The results can be seen in Figure 14 in the last row of images for the first 5 of these cylinder samples. These images were generated from an arbitrary layer within the given part. The CT data was collected by a 225 KeV x-ray CT system with a 9.48 µm voxel size.

These images were generated from an arbitrary layer within the given part. The CT data was collected by a 225 KeV x-ray CT system with a 9.48 µm voxel size. The data was post-processed with ring removal, beam hardening correction, intensity equalization in Z-slices, background removal, and followed by thresholding segmentation to segment pores.1244Data CorrelationsTo correlate the NIR signal intensity to the anomalies detected by the CT data, the NIR data and CT data need to be quantified. A qualitative correlation is present in Figure 14 as the intensity section of data shows an increase in spatter and geometrical representation of the desired part, as well as an increase in intensity.

When laser power, and subsequently NIR signal intensity, increase there is an increase in density in the CT layer images.Figure 14: CT and NIR comparison.To generate a quantitative correlation between the CT data and NIR data, a value was calculated as a representative indication of NIR intensity and CT density. The value calculated for NIR signals is cumulative NIR intensity which is the highest layer-wise cumulative pixel value from the layer analyzed in previous sections. The representative value for CT density was calculated by summing the number of non-empty voxels (these were identified as the pixels of the binary CT results that indicated 1 and not 0) for a designated region in all samples, as the samples all had different numbers of layers. The percentage of non-empty voxels was then calculated based on the sample considered to be fully dense (the highest number of non-empty voxels). The correlation between these two values can be seen in Figure 15.

The correlation between these two values can be seen in Figure 15. It can be seen that at NIR cumulative intensity values above 1500 yield final part densities above 90% while NIR cumulative intensity values below 1500 yield final part densities well below 60%, indicating severe porosity may exist.1245Figure 15: CT and NIR correlationConclusionThrough this research and others, it was noticed that IR cameras are expensive, slow, and yield low resolution. In many other published works, IR data has also been shown as still useful and valuable for additive manufacturing process monitoring. NIR data has been suggested to complement and possibly augment IR data [1]; however, this proved unfruitful for this research with our current system configuration.The utility of NIR signal monitoring in L-PBF for process monitoring remains unresolved.

We have shown that NIR cameras are capable of imaging interlayer temperatures or melt pool geometry, but not concurrently due to the different configurations needed with differing process conditions. The accumulated NIR intensity signals do correlate with global porosity trends, but future work is still required to resolve local porosity correlation.Future work in this area of research aims at exploring geometrical defect detection, rather than global porosity detection. Other areas of exploration are the use of NIR as a powder bed temperature measurement device which would require the calibration of NIR signals to temperature and would allow for the ability to quantify thermal gradients with the merged IR and NIR data.

In-situ monitoring of Ultrasonic Additive Manufacturing (UAM) process is crucial for producing parts suitable for load-bearing structural applications. Due to the nature of UAM process, it is necessary to monitor the entire build as opposed to only the just bonded layer. For this purpose, an ultrasonic transducer is used in this study to perform in-situ nondestructive evaluation (NDE) of the entire build after the addition of each new layer. This has been successfully implemented first on a manually operated research UAM machine and then applied on a fully automated commercial grade UAM machine.

This has been successfully implemented first on a manually operated research UAM machine and then applied on a fully automated commercial grade UAM machine. The practical applications of such in-situ measurements for ensuring defect-free part fabrication through closed-loop control of the UAM process control is shown to be possible from the results of this work.INTRODUCTIONUAM is a layer by layer metal based solid state bonding process [1] which is commonly used in conjunction with a CNC mill and offers advantages in material properties as compared to traditional metal joining and forming technologies [2]. In this process, a high level of ultrasonic energy is utilized to create solid state metallurgical bonding between a thin foil and an existing base. Given the right process parameters, 100% linear weld density can be achieved. CNC machining after depositing each layer is done to achieve required shapes and dimensions, enabling a freeform fabrication.

CNC machining after depositing each layer is done to achieve required shapes and dimensions, enabling a freeform fabrication. The additive-subtractive nature and the low-heat solid state processing give the UAM process certain unique capabilities such as completely enclosed internal cooling channels, smart parts with embedded sensors and metal-matrix composites that cannot be produced otherwise [3-5]. Even dissimilar metals that are not metallurgically compatible can be combined to produce multi-material parts without any embrittlement or cracking problems. For several years, the material systems that could easily be bonded using the UAM process were limited to aluminum alloys due to very high ultrasonic power requirement for other engineering alloy materials. Newer, commercial grade UAM systems overcome this hurdle by using a high power transducer and load1557cell, which makes it possible to build components from titanium, nickel and iron based alloy systems [6].

Figure 1(a) shows a schematic of the UAM process and Figure 1(b) shows the manual research machine used in the current study. The load cell shown in Figure 1(b) applies a constant force while the ultrasonic horn laterally vibrates as it moves forward causing intense plastic deformation at the foil/substrate interface forcing the foil to bond to the substrate. The process of ultrasonic bonding involves breaking up of oxide layers on the surface of the sheet metal, followed by nascent metal contact due to intense plastic deformation. Normally, UAM components are transversely anisotropic along the build direction by nature as they are layered composite structures. The quality of bonding at each interface determines the overall component strength.

The quality of bonding at each interface determines the overall component strength. Since UAM was developed for producing end-use functional parts, it is beneficial to monitor the process online to enable closed loop control of the process.High power Load CellMotion Control20 kHz UltrasonicEDM cut Sonotrode (14 µm roughness)(a)(b)FIGURE 1. (a) Schematic of the Ultrasonic Additive Manufacturing consolidation process and (b) SonicLayer R200 manually operated research machine.MOTIVATIONUAM parts may contain several defects such as partial and kissing bonds (even complete delamination) unless the process parameters are continually changed with increasing build height. This is because layer bonding is a vibration dependent phenomenon. During each welding cycle, the entire part vibrates under a compressive load. This can lead to opening of any kissing bonds or partial bonds into delamination.

During each welding cycle, the entire part vibrates under a compressive load. This can lead to opening of any kissing bonds or partial bonds into delamination. Thus, it is imperative to ensure that not only the current layer is fully consolidated, but also all the previously bonded layers are intact. This is not as critical with other additive processes such as selective laser melting where it is enough to monitor the top surface of the build using optical techniques such as infrared imaging. But in the case of UAM, a layer may be fully bonded initially, but can develop into a delamination due to repeated cyclic loading several layers after it has been bonded. It is thus crucial to have a monitoring mechanism that can see through the entire component. Sometimes, IR cameras are used to monitor the top surface of the UAM build [7] or embedded thermocouples are used to measure the temperature within the part [8].

Sometimes, IR cameras are used to monitor the top surface of the UAM build [7] or embedded thermocouples are used to measure the temperature within the part [8]. Both these approaches, while having their merits, are not intended nor are thought to be suitable for online monitoring. In this study, conventional longitudinal mode broadband ultrasonic contact transducers are utilized to take a single point scan after each layer is deposited.1558METHODIn-Situ Ultrasonic Non-Destructive Evaluation SetupThe sonotrode employed in the current study has a surface roughness of 14 µm and it gives an average surface roughness of around 10 µm on the top of each deposited layer. The first in-situ monitoring system design involved a setup that can scan the part from the top as shown in Figure 2 (a). It was soon realized that performing an ultrasonic scan over the part after each layer was not practical due to unintended interruption to the manufacturing process itself.

It was soon realized that performing an ultrasonic scan over the part after each layer was not practical due to unintended interruption to the manufacturing process itself. Thus, a new system was designed to provide access from below the base plate as shown in Figure 2 (b). This setup allows for the user to have a capability of continuous monitoring for the NDE signals without any interruption during part fabrication. The preliminary setup was designed to be a single point normal incident contact measurement aimed at the middle of the build area. [9] has used a similar setup with access from below to measure laser melted AM components using ultrasonic sensorssensor150 µm foilsAl 6061 BaseUltrasonic Transducer access from below(a)(b)FIGURE 2.

(a) C-Scan immersion acoustic microscopy approach with access from the top and (b) schematic of in- situ setup with access from below the base plate.Factorial Design of ExperimentsFor the current online monitoring study, 150 (cid:1)m Al 6061 H18 foil was chosen. While the processing parameters for Al 6061 have been well established using the Solidica Ultrasonic Consolidation system [10-11], the higher power UAM system provides a wider range of possible process parameters. Two sets of experiments were performed with similar bonding parameters, one on a manually operated research machine and another on a fully automated commercial grade machine. To study the feasibility of in-situ measurements, an ultrasonic NDE setup was first designed and installed on the research machine. The controllable parameters in UAM are applied force (F), velocity of bonding (V) and amplitude of vibration (A). Of these three parameters, the effect of amplitude is most significant.

The controllable parameters in UAM are applied force (F), velocity of bonding (V) and amplitude of vibration (A). Of these three parameters, the effect of amplitude is most significant. Varying all three parameters together requires a large number of samples to statistically separate the coupled effects of the build parameters. Hence, in this study, force and velocity were kept constant and only amplitude was changed. A three replicate amplitude factorial was carried out on the SonicLayer R200 machine at a constant force of 5000 N (1120 pounds) and a constant velocity of 85 mm/s (200 inch per minute). Rectangular specimens with flat top surface measuring 1” x 3” with a built height of 45 layers were manufactured. In each replicate the amplitudes had 6 values ranging from 27 - 47 µm.

Rectangular specimens with flat top surface measuring 1” x 3” with a built height of 45 layers were manufactured. In each replicate the amplitudes had 6 values ranging from 27 - 47 µm. Based on the data gathered from the research machine, the range of amplitudes was readjusted to be 25 - 35 µm for the subsequent measurements on a commercial grade Fabrisonics SonicLayer 7200 UAM machine.1559In-Situ Data AcquisitionA 5 MHz longitudinal mode broadband contact transducer was driven using a JSR Ultrasonics, DPR 500 remote pulser/receiver module. Ultrasonic data was processed by using a 1.5 GHz 14- bit Acquisition Logic PCI digitizer A 500 V negative pulse was used to drive the transducer with a damping of 50 Ω at a pulse repetition rate of 200 Hz. The processed raw ultrasonic data was exported and analyzed in MATLAB. Figure 3 shows A-Scan signals collected on the base (before depositing any layers), after 5 layers, and after 10 layers of UAM builds.

Figure 3 shows A-Scan signals collected on the base (before depositing any layers), after 5 layers, and after 10 layers of UAM builds. It can be observed that the signal starts shifting to the right as the build height increases. It is also clear that the reflections from the interface are significantly small. In this study, measurements were only taken after building each layer. It should be noted that each interface goes through several cycles of load as more layers are built. Of all the interfaces in the build, the very first interface formed between the base and the UAM build is of primary importance. Delamination of the first interface during manufacturing may ruin the part. Due to the nature of the loading condition in the UAM system, the base/build interface tends to be the weakest link as it experiences the maximum shear loads on account of being farthest from the Sonotrode head. Often times the base is used as a part of the final part or component.

Often times the base is used as a part of the final part or component. In such cases, it becomes critical to have a sound base/build interface.Static data gateδ(t5,4)Dynamic data gateStatic data gateδ(t10,4)Dynamic data gateStatic data gateFIGURE 3. Shift in the time of flight of back wall echoes as new layers are added.Data AnalysisA total of 18 test specimens, 45 layers in each, were built. During manufacturing, ultrasonic data was collected one per layer. Analysis of all this data was carried out by importing each of these waveforms into MATLAB.1560Data Acquisition : jth layerStatic data gate aimed at the Interface to check for delaminationStore Data in MATLAB StructureUltrasonic VelocityShift Dynamic data gate by ∆	(cid:3)	= time taken for pulse echo to go through one layer of buildAttenuation of back wall echo for quality of build estimation(j= j+1) UAM layersFIGURE 4.

Logic used for data acquisition and analysis.The two crucial data sets in the A-Scan signals are the interface reflection (the initial position of the peak on the base) and the current top surface echo (i.e., base plus build). To be able to analyze data in the far field, the back wall 3 and 4 echoes from the top of the build were investigated. As a new layer is added, two data gates of information are stored in a MATLAB structure, the first one being the region around where the interface peak is expected (depending on base plate thickness), and the second one being the shift in the data gate of a specific echo so as to capture the essential echo waveform. This shift is different for back wall 3 and back wall 4 due to different travel times. This moving data gate always captures the back wall echoes (i.e. the reflection from the top of the specimen) and the static data gate captures any reflections from the interface as shown in Figure 3.

the reflection from the top of the specimen) and the static data gate captures any reflections from the interface as shown in Figure 3. Storing all the echo waveform data in a MATLAB structure helps easily form correlations between different build parameters and between different layers of the same build.Ultrasonic Velocity and AttenuationLinear ultrasonic parameters that can be readily calculated from a time domain signal are ultrasonic velocity and attenuation. Accurate calculation of ultrasonic velocity requires knowledge of the thickness of the component that is being measured. The correlation between ultrasonic velocity and mechanical properties is well known. The extent of consolidation of a foil depends on the input parameters used. Due to the large amount of plastic deformation and loads applied, the foil thickness reduces after bonding. The average thickness value of each layer is dependent on the process parameters used.

Due to the large amount of plastic deformation and loads applied, the foil thickness reduces after bonding. The average thickness value of each layer is dependent on the process parameters used. In the current study, the thickness of the builds consisting of 45 layers was measured to obtain the average layer thickness for each of the parameter sets.To find the time delay due to the addition of a new layer of UAM build, two approaches can be followed. The first one involves choosing two/three consecutive back wall echoes of a single waveform and finding the group delay between them. The second approach involves choosing a back wall echo in the far field and following the time shift in that back wall echo due to the addition of a new layer. The advantage of the second method is that all measurements are performed relative to the base material which is acting like a buffer rod.

The advantage of the second method is that all measurements are performed relative to the base material which is acting like a buffer rod. Such a buffer zone is helpful to measure the change due to a small thickness addition as shown by Papadakis [12]. This also eliminates the effect of coupling and the phase shift associated with it. Thus, the second approach has been chosen1561as a suitable analysis method for the current study.

The value of δ(tj,N) is calculated from the known average layer thickness for a given parameter set dparam , the back wall echo number N, layer number j, and the velocity C for Al6061. tmax[Wave]	represents the time at which the maximum of the rectified envelope signal occurs, which is used to calculate the group delay Tj,N. Vgroup,j,N is the group velocity through j layers of UAM build. Lj,N is the relative attenuation in dB due to the addition of new UAM layers. Amp(FFTj,N (ftrans)) represents the magnitude of the FFT of the Wavej,N signal calculated at ftrans which is the center frequency of the transducer.

Amp(FFTj,N (ftrans)) represents the magnitude of the FFT of the Wavej,N signal calculated at ftrans which is the center frequency of the transducer. The attenuation loss is calculated as a relative amplitude change of the center frequency component of the Nth back wall echo signal after j layers of build.RESULTSDelamination Detection and Parameter Estimation for the Research UAM Machine Three replicates of a large parameter set were built in the research machine to determine under what conditions the base/build interface starts separating. At a constant force of 5000 N (1120 Pounds) and a constant build velocity of 85 mm/s (200 Inch per minute), the amplitude of UAM builds was varied from 27 - 47 µm on the research machine.

At a constant force of 5000 N (1120 Pounds) and a constant build velocity of 85 mm/s (200 Inch per minute), the amplitude of UAM builds was varied from 27 - 47 µm on the research machine. Figure 5 shows how the static and dynamic data gates work to detect delamination for a typical case of 45 µm vibration amplitude.Until 8 layers of UAM build (1.2 mm height), the signal from the very top of the part seems to be the strongest and it moves in accordance with the dynamic data gate. By layer 10, signals started showing up on the static data gate. By layer 14, there is a clear and complete delamination at the center of the base/build interface which is large enough to reflect the entire signal back. This is also clear from the dynamic data gate as there is no transmission through the UAM build. After 23 layers, the build showed the first signs of visual delamination from the base while indications from online monitoring started as early as layer 10.

After 23 layers, the build showed the first signs of visual delamination from the base while indications from online monitoring started as early as layer 10. It is evident that ultrasonic signals can detect an indication of delamination for the base/build interface much earlier than the damage stage at which delamination can be visually detected. This capability was used to search across a range of parameter sets to identify the amplitudes when the base/build interface bond breaks down hence limiting the ability of the in-situ ultrasonic monitoring system.1562(1)(2)(3)Static data gateStatic data gateDynamic data gateStatic data gateDynamic data gateStatic data gateDynamic data gateStatic data gateDynamic data gateStatic data gateDynamic data gateFIGURE 5. Delamination detection in a build with 45 µm vibration amplitude.

Delamination detection in a build with 45 µm vibration amplitude. By using the static data gate coupled with the dynamic data gate, indications of delamination can be detected.Table 1 shows the results across three replicates of the six factor amplitude factorial (27 - 45 µm) carried out on the research machine. It is observed that the components started delaminating at the base when increasing the amplitude above 35 µm. In many cases, the delamination becomes visually observable several layers after it has been detected by ultrasound. When two out of three replicates delaminate from the base, such parameter sets are considered unsuitable. At a very high amplitude such as 47 µm, the foils start sticking to the sonotrode, marking the upper limit of the amplitude. It is clear that amplitudes above 35 µm are too harsh on the base/build interface for in- situ monitoring.

It is clear that amplitudes above 35 µm are too harsh on the base/build interface for in- situ monitoring. Hence the parameters for building components on the commercial machine were chosen to be in the range of 25 to 35 µm with increments of 2 µm.TABLE 1. Observations of base/build delamination during part fabrication on the research machine. The number in parentheses indicates after how many layers the delamination can be detected.Amplitude27 µm 31 µm 35 µm 39 µm 43 µm 47 µmReplicate 1 Good Good Good Delamination (18) Delamination (19) Delamination (7)Replicate 2 Good Good Delamination (23) Delamination (21) Delamination (14) Delamination (10)Replicate 3 Good Good Good Good Delamination (16) Delamination (12)1563In the components built using the commercial machine, ultrasonic monitoring suggested delamination in two (33 and 35 µm builds) out of eighteen parts although there were no visual indications.

Unlike the manual research machine, with the automatic commercial machine, the trends in ultrasonic velocity and attenuation did not vary widely across the replicates for a given parameter set. Due to material sticking onto the surface of the UAM horn, uniform bond quality is often achieved after a thin layer of the material is formed on the horn which is called a patina. It is found that similar parameters can have different bond qualities if parts are manufactured using a fresh horn or a horn with a patina. After the formation of a good patina, in the third replicate of the factorial design, none of the components delaminated from the base and there were no outliers in the trends.Ultrasonic Group VelocityUltrasonic group delay Ŧj as calculated by Equation 1 is always the cumulative effect of the j layers of UAM build on the delay of the ultrasonic signal. Hence, it gives the average group delay due to the current UAM build.

Hence, it gives the average group delay due to the current UAM build. The layer wise cumulative group delay is expected to increase linearly if one assumes that each layer has uniform bond quality. It has been observed that the group delay is largely dependent on the build parameters. The null hypothesis is to assume that the UAM builds are perfect and there is only a thickness change which causes the group delay. To test this hypothesis, the average thickness after bonding 45 layers has been measured with a vernier caliper. The average layer thickness was found to monotonically decrease from an average of 148 µm for 25 µm build amplitude to an average of 140 µm for 35 µm build amplitude. This is understandable as higher build amplitudes lead to more consolidation and plastic deformation, resulting in greater reduction in foil thickness.

This is understandable as higher build amplitudes lead to more consolidation and plastic deformation, resulting in greater reduction in foil thickness. However, on the calculation of ultrasonic velocity from these thickness values, it was found that the null hypothesis is rejected and the change in group delay is due to imperfect interfaces within the components. Since accurate thickness values are not known for the first few layers, the group velocity calculations in Figure 6 are shown starting from layer 6 onwards. The group velocity values for amplitudes 25 and 27 µm are calculated until the actual signal amplitude from the top of the UAM build drops to a value when it cannot be differentiated from reverberations within the part. This gives some spurious results which can be ignored.

This gives some spurious results which can be ignored. Also, once the component turns so bad that it highly attenuates the ultrasonic signal, it is no longer worth monitoring.As can be seen in Figure 6, the group velocity clearly differentiates between various build amplitudes. It indicates that the 33 and 35 µm builds are clearly the best. The 29 and 31 µm builds are significantly better than the 25 and 27 µm ones. Parts with 25 and 27 µm amplitudes obviously have several imperfect interfaces (confirmed by optical microscopy). It is well known that imperfect interfaces act like a filter and cause a group delay [13-14]. This group delay is observed as a change in apparent ultrasonic velocity. The values of ultrasonic velocity that appear to be greater than that of the base material (6.3 mm/µs) and the decreasing trend of ultrasonic velocities can be attributed to three reasons.

The values of ultrasonic velocity that appear to be greater than that of the base material (6.3 mm/µs) and the decreasing trend of ultrasonic velocities can be attributed to three reasons. The first and most probable being that the exact thickness values are unknown and are estimated from the average thickness after 45 layers of build. The second could be that the acoustoelastic effect is changing the ultrasonic velocity. The third effect could be that as the layer number increases ultrasound travels through more imperfect interfaces thus causing a decreasing quality of build. It is believed that a combination of the above factors is responsible for observed decreasing trends in ultrasonic group velocities. The decreasing trend is not observed in components of lower quality (25, 27 µm). It is hypothesized that the reason for this is the interaction of the ultrasound echoes from the top of the specimen with those from within1564the specimen.

It is hypothesized that the reason for this is the interaction of the ultrasound echoes from the top of the specimen with those from within1564the specimen. These interference effects could distort the measured ultrasonic group velocities especially when the echoes from within the sample are of significant strength. In such cases, calculating the real group velocity pertains to solving the inverse problem of evaluating the part quality given different levels of imperfections within a set of layers. This is beyond the scope of the current work. The interference effects can be further observed through attenuation results.6.56.3] s µ / m m6.1[y t i c o l e V c i n o s a r t lU5.95.75.525 µm27 µm29 µm31 µm5.333 µm35 µm5.110 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 Layer NumberFIGURE 6.

Ultrasonic group velocity as layers added under various amplitude conditionsUltrasonic AttenuationUltrasonic attenuation is shown in Figure 7. It represents a very interesting phenomenon which occurs due to the periodic nature of the UAM builds. Ultrasonic wave propagation through periodic structures is well studied in layered composites [15]. Due to the periodic nature of the structures and the imperfect interfaces at each layer, there is a small reflection from every single interface of the UAM build. These signals interfere constructively and destructively to form periodically changing amplitudes which also attenuate with increasing number of layers. This effect is most prominent in the low amplitude builds (25 and 27 µm). To a lesser extent, the oscillating periodic nature was also observed for build amplitudes of 29, 31 and 33 µm.

This effect is most prominent in the low amplitude builds (25 and 27 µm). To a lesser extent, the oscillating periodic nature was also observed for build amplitudes of 29, 31 and 33 µm. The aforementioned phenomenon coupled with the fact that the attenuation loss through the build is lesser as the amplitude increases, implies that each interface reflection is much smaller when the build amplitude is higher. This can be explained by the fact that as the build amplitude is increased until 35 µm the interface quality increases and the UAM layers start behaving like solid aluminum. It is also evident that a bad interface would lead to a higher reflection which would lead to a constructive or destructive interference with the waveforms from the top of the build as subsequent layers are built up.

Thus higher attenuation loss and larger oscillations in their relative amplitude1565imply bad interfaces while lower attenuation loss and smaller oscillations indicate better bond quality.0510]B d [15n o i t a u n e t t2025 µmA27 µm2529 µm31 µm3033 µm35 µm350246810 12 14 16 18 20 22 24 26 28 30 32 34 36 38 40 42 44 46Layer NumberFIGURE 7. Relative ultrasonic attenuation of the 3rd back wall echo signals due to layer build upCONCLUSIONS AND FUTURE WORKAn in-situ ultrasonic non-destructive evaluation system has been developed to monitor bond quality of a build produced by Ultrasonic Additive Manufacturing. An online ultrasonic measurement approach proved to be effective for detecting the inception of small delamination at the base/build interface. This information could be used to adjust the build parameters online to save a component from failure during build.

This information could be used to adjust the build parameters online to save a component from failure during build. Results from the current in-situ NDE study indicate that linear ultrasonic parameters such as ultrasonic group velocity and attenuation are sensitive to the variations in build amplitude and interface bond quality. Good quality builds exhibited higher ultrasonic velocity, lower attenuation and smaller oscillations in the attenuation values. Fabricated parts with imperfect interface bond conditions lead to lower sound velocities due to signal delay and higher attenuation values because of absorption/scattering through poorly bonded interfaces.

The current work shows a qualitative correlation between ultrasonic signatures and UAM build parameters and lays the foundation for a potentially useful in-situ monitoring tool for online closed loop process control.ACKNOWLEDGEMENTSThe authors would like to thank Mark Norfolk for access to the Soniclayer R7200 commercial system as well as Hilary Johnson and Brian from Fabrisonics for helping out with the experiments. Joe Vickers and Sudeep Vickram made significant contributions towards work done on the Soniclayer R200 machine. The authors would also like to thank Dr. Peter B Nagy for his valuable advice. Finally, we would like to acknowledge the source of funding from the Office of Naval1566Research Grant ONR-BAA #14-004-1110689, Cyber Enabled Manufacturing Systems, Acoustic resonance techniques for online certification and offline qualification of metal based additive manufacturing technologies.REFERENCES 1. White, D., & Carmein, D. E. (2002). U.S. Patent No. 6,463,349.

The spectral data can be used to calculate the temperature of the vapor plume and correlated with the melt-pool size. The effects of varying the atmosphere and pressure on the OES signal are also explored. These results demonstrate that OES can provide useful feedback to the SLM process for process monitoring and part validation. The challenges implementing OES in-line on a commercial SLM platform are discussed. This work was funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy.1. IntroductionSelective Laser Melting (SLM) is a powder-bed based additive manufacturing (AM) process in which 3-Dimensional parts are built layer-by-layer. One challenge in AM processes are that the temperature history varies with the geometry. Measuring local processing conditions is critical for part validation and provides opportunities for feedback-based control.

Measuring local processing conditions is critical for part validation and provides opportunities for feedback-based control. While photodiodes provide some information about the state of the process, the spectral content of light emitted from the melt pool and plume gives potential insight into the temperature, chemistry and pressure surrounding of the melt pool.Optical Emission Spectroscopy (OES) has been demonstrated with laser welding to provide information about the chemical species and calculate the temperature and electron density in the vapor plume [1-3]. The spectroscopic data has also been correlated with features in the welds such as the depth to width ratio. OES methods developed for laser welding have also been extended to monitor powder metal deposition Direct Energy Deposition (DED) blown-powder process.

OES methods developed for laser welding have also been extended to monitor powder metal deposition Direct Energy Deposition (DED) blown-powder process. Song and Mazumder measured chromium emission in directed metal deposition and used line intensity ratios to calculate plasma temperatures and used emission signal to predict chromium composition real-time during processing with H13 tool steel [4]. Ya et al. demonstrated the ability to relate OES signals to clad quality in laser metal deposition [5]. Nassar et al.

Ya et al. demonstrated the ability to relate OES signals to clad quality in laser metal deposition [5]. Nassar et al. used OES as a process monitoring tool to correlate emission line-to-continuum ratios to lack of fusion defects during directed-energy deposition of Ti-6Al-4V [6].Work performed and funded by The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing & Technologies, LLC under contract number DE-NA0002839.2192While OES has been used in blown-powder AM, there are limited reports of applying OES to powder bed based AM. Dunbar et al. mounted a spectrometer at a fixed location in a powder bed based 3DSystems Prox-200 build chamber and reported emission signal as a function of defocusing the process laser [7].

Dunbar et al. mounted a spectrometer at a fixed location in a powder bed based 3DSystems Prox-200 build chamber and reported emission signal as a function of defocusing the process laser [7]. Dunbar and Nassar used photodiodes with band pass filters focused off axis in the build chamber to monitor the chromium emission line to continuum ratio during processing of Inconel 718 with a 3DSystems Prox-200 powder bed AM system [8].The galvo-scanner steered laser beam and moving melt pool in the global reference frame complicates implementation of OES for the SLM process. Inserting the spectrometer optics into the beam path allows the collection volume to also be guided by the galvo-scanner. Although the solid-angle is limited, the approach still allows species and plume temperatures to be determined. This paper demonstrates OES of the SLM process with 304L stainless steel.2.

Although the solid-angle is limited, the approach still allows species and plume temperatures to be determined. This paper demonstrates OES of the SLM process with 304L stainless steel.2. Experimental ApproachExperiments to evaluate OES of SLM were completed using a home-built SLM system and an Andor Technology SR-750 spectrometer. The SLM system used an IPG Photonics YLR-500 continuous wave fiber laser (λ = 1070 nm) with an IPG D30 collimator. The x-y position of the laser beam was controlled by a SCANLAB hurrySCAN and the laser was focused by a 340 mm focal length f-θ lens (measured 1/e2 beam diameter of 145 μm). Data reported in the following section was collected using a 600 l/mm diffraction grating installed in the spectrograph and an Andor Technology iStar 734 series ICCD. The spectrometer was coupled to the SLM system as illustrated in Fig.

The spectrometer was coupled to the SLM system as illustrated in Fig. 1 using a DMLP Dichroic mirror (Thorlabs).Galvo- scannerDichroic MirrorFiber Laser Collimatorf-θ LensBuild ChamberMirrorΩ = 2 mSrSpectrometer CollimatorFiber Laser Optical SignalFIG 1: Schematic of SLM system optical components with spectrometer inserted into the beam path.Relative intensities for optical emission lines are described by Eq. 1 [4]:Imn∝⋅ N A mmn⋅⋅ h νmnwhere Imn is the intensity, Nm is the upper state population, Amn is the transition probability, h is the Planck constant, and νnm is the frequency. Emission lines detected during SLM of 304L in this work correspond to neutral chromium and iron (determined species from NIST database). Figure 2 contains representative emission signal measured during laser melting a 304L stainless steel2193(1)powder bed with a laser power of 300 W and scan speed of 675 mm/s. In Fig.

Figure 2 contains representative emission signal measured during laser melting a 304L stainless steel2193(1)powder bed with a laser power of 300 W and scan speed of 675 mm/s. In Fig. 2, both wavelength ranges investigated contain strong chromium emission lines and weaker iron lines.120050001000(a)Cr CrCr4000(b)Cr] . u . a [y t i s n e t n I800600400FeFeFe CrCrFe] . u . a [y t i s n e t n I30002000CrFeCrCrCrCr20010000415420425430Wavelength [nm]4354400520525535 Wavelength [nm]530540FIG 2: Average optical emission signal collected during processing powder 304L in argon build chamber atmosphere.Figure 3 is an example of using OES signals to differentiate materials. The signals were collected during laser melting of solid plain carbon steel and solid 304 stainless steel samples with 300 W.

The signals were collected during laser melting of solid plain carbon steel and solid 304 stainless steel samples with 300 W. Figure 3 shows both the plain carbon and stainless steel samples contain emission lines for iron and manganese. The signals in Figs. 2 and 3 give important information about the composition of the material during processing. The species identified in the vapor plume through OES should be consistent for a particular material. This means OES can be used to flag contaminated powder beds if unexpected emission lines are present in the signal.] . u . a [1.00.8(a)Cr Cr Cr] . u.

This means OES can be used to flag contaminated powder beds if unexpected emission lines are present in the signal.] . u . a [1.00.8(a)Cr Cr Cr] . u . a [1.00.8(b)y t i s n e t n Ie v i t a e Rl0.60.40.2Plain Carbon Steel 304 StainlessMnFey t i s n e t n Ie v i t a e Rl0.60.40.2Plain Carbon Steel 304 StainlessCrCr0.0405410420 Wavelength [nm]4154254300.0520525530535Wavelength [nm]540FIG 3: Differentiation of plain carbon steel and stainless steel through optical emission signals.3. Results and DiscussionIn addition to differentiation of materials during processing, OES can also be used to determine plume temperatures. Emission signals reported in the following figures are time series data averaged over laser melting 5×5 mm2 areas of a 50 µm thick 304L stainless steel powder layer.

Emission signals reported in the following figures are time series data averaged over laser melting 5×5 mm2 areas of a 50 µm thick 304L stainless steel powder layer. Single layer samples were produced by processing the powder bed with varied laser powers and build chamber conditions (atmosphere and pressure). As shown in Fig. 2, the average signals2194545generated over a layer contained an incandescent background with optical emission signals superimposed. The OES data for this section was processed by excluding the optical emission lines to fit the background to a fourth-order polynomial. This fit was then subtracted from the average data including optical emission signals to remove effects of the incandescent background.

This fit was then subtracted from the average data including optical emission signals to remove effects of the incandescent background. The processed OES signals were then used to calculate the plume temperature and correlated with sample melt pool size.Figure 4 (a) is the average optical emission spectra for 304L as a function of laser power collected during processing in an argon atmosphere. Assuming local thermodynamic equilibrium, plume temperatures can be calculated by comparing the intensities of two emission lines of the same species with−> ⋅. Temperature is calculated with OES data by using Eq. 2 [9]:E 2E 1k TT=k⋅ln( ) E 1  I A g 1 2  I A g  2 1 1−E 2 ⋅ ⋅⋅ ⋅2⋅ λ 1 ⋅ λ 2  where T is the temperature, E is the energy, k is the Boltzmann constant, I is emission line intensity, g is the statistical weight, and λ is the wavelength.

The chromium emission lines at λ1 = 529.827 nm and λ2 = 532.834 nm satisfy all conditions and have previously been used in literature to calculate the plume temperature [5]. The intensities (I1 and I2) for λ1 and λ2 were determined as the intensity above background using the method described in [10].

The intensities (I1 and I2) for λ1 and λ2 were determined as the intensity above background using the method described in [10]. Figure 4 (b) is the average plume temperature for processing 304L in an argon atmosphere with varied laser power.Table 1: Constants for selected chromium lines used to calculate the plume temperature.λ [nm]1 Amn [s]gm1 Em [cm ]529.8273,300,000526,797532.83462,000,0001142,26310006000800Cr(a)5500] .u.a[y t i s n etnI6004002000200 W 250 W 300 W 350 W 400 WCrFeCrCrCr]K[e r u t a r e p m e T50004500400035003000Chromium Vaporization Point520 522 524 526 528 530 532 534 536Wavelength [nm]2500200250300 Power [W]350FIG 4: (a) Average optical emission spectra and (b) plume temperatures of 304L stainless steel for varied laser power.2195(b)400(2)In Fig. 4, the overall intensities of emission lines increase with laser power.

4, the overall intensities of emission lines increase with laser power. The increased emission signal with laser power directly correlates to higher average plume temperatures during SLM of 304L.In addition to varying with power, OES signals also change with build chamber atmosphere type and pressure. Figure 5 (a) is the average OES signal for processing 304L in air, argon, nitrogen, and low vacuum. The optical emission signal for air is much stronger than the other atmospheres. This could be due to increased material vaporization, which would be reflected in a keyhole mode like appearance of the melt pool. However, micrographs of the melt pool cross- sections (Fig 5 (d-g)) show conduction mode melting dominates regardless of chamber atmosphere. The increased OES signal strength for air could also be explained due to oxidation of the vaporized chromium and iron.

The increased OES signal strength for air could also be explained due to oxidation of the vaporized chromium and iron. The exothermic oxidation process adds heat to the plume increasing the temperature and resulting signal measured through OES [10]. The OES signals for processing in argon, nitrogen, and low vacuum are not amplified from the oxidation process.Figure 5 (b) shows the dependence of optical emission signal on chamber pressure. OES data was collected while increasing the pressure of an argon atmosphere. The intensity of emission lines increases from low vacuum to above atmospheric pressure (0.2 to 800 Torr). The weaker optical emission signals measured near vacuum are possibly due to lower excitation efficiency because of less energy transferring collisions. Figure 5 (c) is the intensity of chromium emission λ=520.6 nm as a function of chamber pressure and also shows the dependence of layer surface quality for selected pressures (100, 400, and 700 Torr).

Figure 5 (c) is the intensity of chromium emission λ=520.6 nm as a function of chamber pressure and also shows the dependence of layer surface quality for selected pressures (100, 400, and 700 Torr). Lower chamber pressures resulted in better surface finishes when compared to layers processed with higher pressure. This means the layers with higher surface finish quality correspond to weaker intensities of optical emission signals.219610001000] . u . a [800600Air Ar N2 Vacuum(a)] . u . a [800600Pressure [Torr] 0.2 100 200 300 400 500 600 700 800y t i s n e t n I400200y t i s n e t n I400200001000800520(c)522524 528 526 Wavelength [nm]2 mm530532P = 300 W v = 675 mm/s (d)Air520522524 528 526 Wavelength [nm](d)N2530] . u.

u . a [600y t i s n e t n I400200Ar λ = 520.6 nm(e)Ar(e)Vac.00200400 Pressure [Torr]60080025 µmFIG 5: OES signal collected during SLM processing with different chamber atmospheres and pressures and micrographs of 304L stainless steel single layers.Figure 6 is the correlation of the average plume temperatures determined by OES to the average melt pool size across single layer samples processed with varied laser power. Larger plume temperatures directly correlate to larger melt pools. Since OES data relates to sample properties, signals can potentially be used as feedback for control of SLM. This has been demonstrated in laser welding where the plume temperature determined by OES was used as feedback to control the penetration depth of the weld [11].

This has been demonstrated in laser welding where the plume temperature determined by OES was used as feedback to control the penetration depth of the weld [11]. Similar OES feedback methods can be applied to SLM for layer-to-layer and point-to-point control.150]m µ[e z Si125(a)Half Width(b)Half Widthlo o P100t l e M e g a r e v A7550DepthD e p t h35004500 Average Plume Temperature [K]40005000550025 µmFIG 6: Correlation of melt pool size to average plume temperatures during single layer processing.2197(b)Ar532(f)(g)4. Challenges Implementing OES with Commercial SystemThe results in the previous sections show the potential value of using OES to monitor the SLM process based on results from a home-built system used for fundamental research and manufacturing of small scale samples.

Implementing OES on a commercial system will allow an investigation of OES signals with correlation to engineering properties determined from larger scale samples. As a part of the first steps to accomplish this, an optical configuration similar to the schematic in Fig. 1 was applied to insert the spectrometer into the beam path of a Renishaw AM250. Figure 7 is an image of the spectrometer optics in-line on the AM250.Fiber Laser Optical SignalVisible CameraMirrorVisible SpectrometerFlip Mount MirrorGalvo- scannerDichroic MirrorLaser Focusing OpticFIG 7: Image of experimental setup for inserting spectrometer optics into laser beam path of Renishaw AM250The experimental setup in Fig. 7 was used in a preliminary investigation to identify challenges in performing OES with the spectrometer optics inserted into the beam path on the AM250.

7 was used in a preliminary investigation to identify challenges in performing OES with the spectrometer optics inserted into the beam path on the AM250. Figure 8 (a) contains the average optical emission signals collected during processing a 10 by 10 mm2 area of a 50 µm thick powder layer of 304L stainless steel in argon. There are no apparent emission lines in Fig. 8 (a). This is due to low transmission of the AM250 galvo-scanner optics from 400 nm to 530 nm. However, it is known from Fig. 5 (a) optical emission signals are stronger for processing in air. The Renishaw AM250 process laser was used to scan and melt solid 304L in air to observe if optical emission signals can be detected. Figure 8 (b) is the successful measurement of chromium emission during laser melting of solid 304L in air with the spectrometer optics inserted into the Renishaw beam path.219825002500] . u . a [20001500Power [W] 150 175 200(a)] . u.

u . a [20001500Power [W] 150 175 200(a)] . u . a [20001500Power [W] 150 175 200Cr(b)y t i s n e t n I1000y t i s n e t n I10005005000 510515520 Wavelength [nm]5255300 510515520 Wavelength [nm]525FIG 8: Signals collected during (a) processing powder 304L in an Argon atmosphere and (b) laser melting solid 304L in air with Renishaw AM250.Figure 9 contains OES signals for two wavelength ranges measured during further investigation with laser melting solid 304L in air with the AM250 process laser (200 W). There are strong neutral chromium and iron signals present in Fig. 9 and the results are comparable to the signals in Fig. 2 obtained during SLM of powder 304L in argon with the home-built system.40006000Cr(a)5000(b)] . u . a [y t i s n e t n I300020001000FeFeCrCrFeCrFeFeCrCr] . u.

u . a [y t i s n e t n I300020001000FeFeCrCrFeCrFeFeCrCr] . u . a [y t i s n e t n I4000300020001000CrCrCr CrCr0415420425430Wavelength [nm]4354400520525535 Wavelength [nm]530540FIG 9: OES signal collected during laser melting solid 304L in air with Renishaw AM250.Figures 8 and 9 show the methods used to insert the spectrometer optics into the laser beam path on the home-built SLM system can be expanded to a commercial platform. However, the current transmission of the Renishaw AM250 galvo-scanner optics and spectrometer optical elements attenuates the emission signal during processing powder 304L with nominal conditions. The challenge of low transmission will be addressed in future work in order to correlate optical emission data with engineering properties of specimens manufactured with the Renishaw AM250.5. Summary and ConclusionsThis paper demonstrated measurement of the SLM plume using in-situ OES.

Summary and ConclusionsThis paper demonstrated measurement of the SLM plume using in-situ OES. The temperature of the plume generated during SLM was determined by comparing ratios of optical emission signal intensities. The calculated plume temperatures correlate well with the melt pool2199530545size. The OES results in this paper also show the dependence of SLM on build chamber atmosphere type and pressure, with higher chamber pressures resulting in stronger intensities of optical emission signals. The correlation of signals measured with OES to sample properties leads to the potential for feedback control with faster spectrometer data. Future work will involve expanding the process windows to include scan speeds and building multi-layer parts with a commercial SLM system to demonstrate the correlation of engineering properties to information determined by using OES.6.

AcknowledgementThis work was funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.7. References[1] Collur M, Debroy T. Emission Spectroscopy of Plasma During Laser Welding of AISI 201Stainless Steel. Metallurgical Transactions B. 1989; 20B:277-86.[2] Szymanski Z, Kurzyna J, Kalita W. The spectroscopy of the plasma plume induced during laser welding of stainless steel and titanium. Journal of Physics D: Applied Physics. 1997; 30:3153-62.[3] Kawahito Y, Matsumoto N, Mizutani M, Katayama S.

Solid Freeform Fabrication Symposium; 2014.[7] Dunbar A, Nassar A, Reutzel E, Blecher L. A real-time communication architecture for metal powder bed fusion additive manufacturing. Solid Freeform Fabrication Symposium; 2016.[8] Dunbar A, Nassar A. Assessment of optical emission analysis for in- process monitoring of powder bed fusion additive manufacturing. Virtual and Physical Prototyping. 2018; 13:14-9.2200[9] Ancona A, Spagnolo V, Lugara P, Ferrara M. Optical sensor for real-time monitoring of CO2laser welding process. Applied Optics. 2001; 40:6019-25.[10] Shin J, Mazumder J. Composition monitoring using plasma diagnostics during direct metaldeposition (DMD) process. Optics and Laser Technology. 2018; 106:40-6.[11] Sibillano T, Rizzi D, Mezzapesa F, Lugara P, Konuk A, Aarts R, Veld B, Anoca A. Closed Loop Control of Penetration Depth during CO2 Laser Lap Welding Process. Sensors. 2012; 12:11077-90.2201.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 2(cid:26)th Annual International(cid:3) Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference(cid:3) Reviewed PaperIn-Situ Preheating in Hybrid Layered Manufacturing for Tooling ElementsFisseha Legesse*, Sajan Kapil*, Rimpy Chabra*, Arun Sharma* and K. P. Karunakaran*Rapid Manufacturing Lab, Mechanical Engineering Department, Indian Institute of Technology Bombay, IndiaAbstractSolidification cracking of hard materials such as H13 tool steel is one of the major problems in metal based Additive Manufacturing (AM) processes. Hybrid Layered Manufacturing (HLM) is one of the metal based AM process which uses Metal Inert Gas (MIG) cladding for addition and CNC milling for subtraction of material. In this work, an in-situ induction heating based preheating system has been developed to solve the solidification cracking problem.

In this work, an in-situ induction heating based preheating system has been developed to solve the solidification cracking problem. The Tungsten Inert Gas (TIG) and Induction heating methods are compared and it has been found that the induction based preheating system can produce better microstructure and sound products. In the experimental procedure, before deposition of a layer the prebuild layer is preheated up to 350-5000C. Also the effect of in-situ preheating on the microstructure of the deposited layers have been studied using Scanning Electron Microscope (SEM).IntroductionIn the last three decades Additive Manufacturing (AM) is playing a great role in the industries in changing 3D CAD models directly to products and prototypes without the use of tooling. Researchers started developing metal AM using laser also called Selective Laser Sintering (SLS) to selectively melt metal powder [1].

Researchers started developing metal AM using laser also called Selective Laser Sintering (SLS) to selectively melt metal powder [1]. Based on material input AM systems for metals nowadays are divided in to three main categories: (i) Powder bed system, (ii) powder feed system and (iii) wire feed system [2]. The comparative study of wire and powder feed metallic AM processes has been investigated by microstructure analysis of the deposited product [3]. The comparison study shows that the microstructure of both the samples are similar, however porosity have been found in the samples produced by powder based AM systems.Wire-feed AM is divided into three groups based on energy sources, these are; laser, arc, and electron beam. Initially, a single bead of material is deposited to make a layer and upon subsequent passes is built upon to develop a 3D object. Wire feed AM are well suited for high deposition rate processing and have large build volumes.

Wire feed AM are well suited for high deposition rate processing and have large build volumes. Deposition rate of laser or electron beam (EB) is of the order of 2–10g/min, whereas deposition rates of 50–100 g/min have been reported in arc- based AM [4, 5]. The fabricated product usually requires more machining between layers and after finish deposition. Therefore, to attain better dimensional accuracy and surface quality for building metallic dies and molds, it was proposed to develop a unique methodology with a hybrid approach of an additive and subtractive processes. It consists of Gas Metal Arc Welding (GMAW) and CNC milling machining respectively [6].One of the applications of AM of metals is in the area of Rapid Tooling (RT). RT methods allow injection molding and die-casting inserts to be built directly from 3D CAD438models.

RT methods allow injection molding and die-casting inserts to be built directly from 3D CAD438models. A computer aided system has been developed for rapid tooling process selection and evaluation of mold design for its manufacturability and cost effectiveness [7]. Arc weld deposition method through hybrid approach of addition of metal through GMAW deposition and subtraction using CNC milling was used to produce a tooling with mild steel material [8].Most of the molds and dies are made out of tool steel materials and hence in this research H13 tools steel material is used to manufacture a tooling element and to see the details of the microstructure. Deposition of H13 tool steel in a wire form is an essential research area for the advantage of getting higher deposition rate.

Deposition of H13 tool steel in a wire form is an essential research area for the advantage of getting higher deposition rate. One of the drawback of tool steel welding or deposition is the issue of solidification cracking.Pre-heating is often used to reduce the residual stresses and the risk of thermal distortion and cracking for tool steel cladding [9]. A research work has shown on 12% Cr steel welding the effect of preheating process on microstructure and to avoid cracking by martensitic microstructure [10]. To overcome this problem in this research induction preheating system is used and its effects are studied in detail.Experiment SetupIn this study a hybrid approach of AM process is used, where Arc weld deposition, generally called MIG/TIG cladding is used as an energy source to deposit a material in a wire form and CNC machine for intermediate machining and final surface finish, figure 1.

This alternate use of additive and subtractive process makes this metallic AM process a hybrid process and it is called Hybrid Layered Manufacturing (HLM). This setup contains a 3-axes milling machine and a MIG/TIG cladding gun that is vertically attached to the spindle housing. A retrofitting of synergetic MIG/TIG cladding on 3-axes milling machine is required in order to carry out the cladding process efficiently [11].(a) Cladding or deposition Unit(b) CNC machining unit(c) Total Setup of MIG Cladding and machining Figure 1 MIG/TIG cladding deposition unit and CNC machining or HLM setup at IIT Bombay439The first essential step in this process is to determine the optimum parameters of deposition. These parameters are welding current, standoff distance and the torch feed rate. The optimum process parameters were validated with the help of Taguchi design of experiments (DOE). Based on the orthogonal array nine different single weld bead depositions was done.

The optimum process parameters were validated with the help of Taguchi design of experiments (DOE). Based on the orthogonal array nine different single weld bead depositions was done. In this experiment H13 tool steel substrate was used to see the overall effect. The composition of H13 tool steel material is shown in table 1. The metal plate was sectioned in to two pieces, perpendicular to the weld bead, by using wire EDM. The finishing of the surface was done by filling, polishing and using a 2% Nital etchant. The profile of the weld bead measurement such as weld width, height and penetration was taken using Stereo Microscope device. The details of the sample reading is shown in the figure 2.

The profile of the weld bead measurement such as weld width, height and penetration was taken using Stereo Microscope device. The details of the sample reading is shown in the figure 2. After taking the result the measured value was calculated using ANOVA method and the optimized MIG cladding parameters were selected.Table 1 Composition of H13 tool steel materialC 0.38Mn 0.39Si 0.94Cr 4.96Mo 1.66V 0.97P 0.019S 0.018Fe BalanceTool steels in the cold and hardened condition always should be preheated for welding or deposition to a temperature not to exceed the tempering temperature and should be maintained as closely as possible at this temperature during deposition operation. The material to be welded must be preheated slowly and uniformly to avoid uneven expansion through localized or partial heating which can lead to cracking.

The material to be welded must be preheated slowly and uniformly to avoid uneven expansion through localized or partial heating which can lead to cracking. It is important to maintain the preheat temperature as constant as possible during welding [12].In the development of the tooling using MIG cladding two types of preheating methods were used. These are TIG arc and Induction heating methods. There are few reasons preheating is essential for MIG cladding of H13 tool steel material. The first one is to slow down the cooling rate and second allows the back filling of molten metal to be fed in the proper dendritic position so that solidification cracking can be avoided [13].(a) Single Weld bead Deposition based on DOE440(b) Nomenclature of single weld bead(c) An image from the stereo microscopeFigure 2 H13 Tool steel single bead deposition experimentIn the first few layer of MIG cladding of H13 material, TIG arc was used for preheating the substrate.

This preheating system was selected based on the availability of the setup, figure 3a. After few layers of MIG cladding it is observed that the cracking started at the point of the neck of the part, see figure 3b and 3c. This problem might be happened because of two reasons. The first one is the insufficient retention of preheating temperature before MIG cladding. The second reason might be oxidation of the substrate material due to TIG arc heating. This claim is clearly seen on figure 3a. This type of oxidized surface was observed while building a part for the few layers.(a) TIG arc preheating(b) Few layers of MIG cladding(c) After machiningFigure 3 TIG arc preheating unit and Cracked H13 MIG cladding objectIn the second attempt the experiment was done using induction preheating method. In this method the substrate was heated up to 4500c and above before each layer of MIG cladding.

In this method the substrate was heated up to 4500c and above before each layer of MIG cladding. The induction heating device was designed for cladding purpose to get an output of 3.2KW power and frequency range of 10 - 20 kHz. In this preheating method the required temperature was achieved in short period of time and relatively less oxidation was seen. The shape of the heating coil is called pancake type as shown in the figure 4.Figure 4 Induction preheating unitThe 3500c and above preheating temperature maintained while layer by layer deposition was done. In the development of the tooling element using H13 tool steel material deposition441the geometry of the part to be deposited was selected based on the demand from the tooling industry.Result and DiscussionIn this research to see the effect of induction preheating on MIG cladding of H13 tool steel material, two samples were prepared.

One sample is using induction preheating of temperature above 3500c and the other is without preheating. The deposition parameters taken for the for the MIG cladding are current 210A, stand of distance 10mm and torch speed 1800 mm/min. The deposited near net shape and finished machined block are shown in figure 5a and 5b.(a) Near net shape Figure 5 H13 tool steel samples for microstructure test with and without Induction preheating(b) Finished machiningThe micro hardness of the samples in the coordinate directions were tested. The deposition direction was along x-axis, the successive weld bead or pitch direction was along y- axis and the block was completed adding layer by layer in the z-direction. The part was finished machined and removed out from the substrate material by EDM wire cut. In figure 6a shows two machined blocks deposited using in-situ preheating process and without preheating respectively.

In figure 6a shows two machined blocks deposited using in-situ preheating process and without preheating respectively. Micro hardness machine was used to measure the hardness in the coordinate directions. The result is shown in figure 6b, 6c & 6d. The hardness of the deposited H13 tool steel block without preheating shows more hardness than the part produced by In-situ preheating system.

The result is shown in figure 6b, 6c & 6d. The hardness of the deposited H13 tool steel block without preheating shows more hardness than the part produced by In-situ preheating system. The trend looks similar in all coordinate directions.70Preheating Without preheating65s s e n d r a H6055500515 10 X coordinate20(a) Samples for hardness test(b) Harness variation along x direction442257050Preheating Without preheating45Preheating Without preheating6540s s e n d r a H60s s e n d r a H3530552520500510 Y coordinate15150124 Z coordinate356(c) Hardness variation along y direction(d) Hardness variation along z directionFigure 6 Hardness comparision along the coordinate axes directionThe required specimen was prepared by cutting out a size 10 X 10mm from the center of the deposited sample using EDM wire cut.

The samples were polished using a polishing machine and then socked in an etchant solution of 2% Nithal for about one minute. The microstructure of both samples were seen using electron microscope (Axiocam iCc3, ZEISS). The picture of the two sample results with induction preheating and without preheating are shown in figure 7.(a) 5x magnification(d) 5x magnification(b) 10x(e) 10x magnification443(c) 20x magnification Figure 7 a, b, & c are microstructure pictures without preheating and d, e, & f are with In-situ preheating(f) 20x magnificationThe pictures taken using electron microscope clearly shows the difference between the two methods. Figure 7a is a part produced without preheating, shows more number of porosity and figure 7d a part produced with in-situ preheating, shows few porosities.

Figure 7a is a part produced without preheating, shows more number of porosity and figure 7d a part produced with in-situ preheating, shows few porosities. Figure 7e & 7f are the pictures of the part deposited using in-situ preheating shows clear elongated grain structure and dendritic formation, whereas figure 7b & 7c deposited without preheating the pictures are shown finer grain structure arraignment. This might be one of the reason the H13 tool steel produced with MIG cladding without preheating are shown more hardness.Another issue in H13 tool steel MIG cladding is the solidification cracking phenomena. This effect is observed using Scanning Electron Microscope (SEM) (EVO/18 ZEISS). Figure 8a shows that the porosity is the starting point of solidification cracking in the deposition process without preheating. Figure 8b shows the number of porosity in a specific area of the sample.

Figure 8b shows the number of porosity in a specific area of the sample. The microstructure formation of the part looks good at specific area, figure 8c.(a) Porosity observed with SEM and the formation of solidification cracking444(b) Porosity seen at the magnification of 119x Figure 8 The microstructure and porosity of H13 MIG cladding without preheating(c) Microstructure seen at 5.27kx magnificationThe SEM reading of the sample of H13 tool steel material deposited with In-situ preheating shows no solidification crack formation around the porosity. This phenomenon is clearly shown in the figure 9a.

This phenomenon is clearly shown in the figure 9a. The microstructure shows a better dendritic formation and the number of porosity observed few compared to the part produced without preheating, figure 9b & 9c.(a) Porosity observed using SEM without formation of solidification cracking(b) Porosity seen at the magnification of 119x Figure 9 The microstructure and porosity of H13 MIG cladding with In-situ preheating(c) Microstructure seen at 5.27kx magnification445ConclusionIn this work, an induction based in-situ preheat procedure has been developed for metal based Additive Manufacturing (AM) processes. This preheating process has been studied for the deposition of H-13 tool steel. The hardness tests are carried out and microstructure analysis has been done. It is observed that preheating can reduce the hardness of the final component in all three directions.

The hardness tests are carried out and microstructure analysis has been done. It is observed that preheating can reduce the hardness of the final component in all three directions. The results obtained by electron microscopy clearly shows that the number of porosities are more in case of non in-situ preheating. Also the growth of grains are more finer in this case which may further lead to the crack formation. The cracks and porosities are more clearly visible in the case of non-pre-heated samples using Scanning Electron Microscopy (SEM). While the sample made by using in-situ heating have less porosity and no crack related issues. Hence, induction based in-situ preheating is recommended for the manufacturing of H13 tool steel components by AM processes through MIG cladding process.446References1. D. L. Bourell, H. L. Marcus, J. W. Barlow and J. J. Beaman: ‘Selective laser sintering of metals and ceramics’, Int. J. Powder Metall., 1992, 28, 369–381.2. William E.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIN-SITU PROCESS MONITORING AND EX-SITU PART QUALITY ASSESSMENT OF SELECTIVE LASER SINTERING USING OPTICAL COHERENCE TOMOGRAPHYAdam Lewis1, Michael Gardner2, Austin McElroy2, Thomas Milner2, Scott Fish3, Joseph Beaman31McKetta Department of Chemical Engineering 2Department of Biomedical Engineering 3Department of Mechanical Engineering Cockrell School of Engineering, The University of Texas at Austin, Austin, TX 78705AbstractWidespread commercial adoption of Selective Laser Sintering has been hindered by inadequate quality and consistency of manufactured parts. Improved process monitoring and control have the potential to improve part quality and thus increase adoption of SLS for various applications.

Improved process monitoring and control have the potential to improve part quality and thus increase adoption of SLS for various applications. In this paper, optical coherence tomography (OCT) is explored as a new process monitoring tool in SLS polymer printing. The basic operating principles behind OCT are reviewed to illustrate the potential monitoring capabilities followed by results for both in-situ process monitoring and ex-situ examinations of built parts comprised of various polymers. Capabilities and limitations of OCT in each application are discussed.IntroductionThe Additive Manufacturing (AM) industry continues to expand rapidly, growing by over US $1 billion for the 2nd year in a row with a 2016 growth rate (CAGR) of 25.9% [1]. Selective Laser Sintering (SLS) is one of the earliest AM techniques and continues to gain further use in prototyping and end-use applications in the medical and aviation industries among others [2].

Selective Laser Sintering (SLS) is one of the earliest AM techniques and continues to gain further use in prototyping and end-use applications in the medical and aviation industries among others [2]. However, SLS manufactured parts continue to be plagued with irregular surface quality and inconsistent strength and dimensions among other issues. In large part, these problems can be attributed to the lack of process monitoring tools in SLS machines. The extent of monitoring in most commercial SLS machines consists of data acquired from a few thermocouples and pyrometers during operation. New and/or improved process monitoring tools coupled with control strategies can help reduce many of the part quality issues mentioned.

New and/or improved process monitoring tools coupled with control strategies can help reduce many of the part quality issues mentioned. In this paper Optical coherence tomography (OCT) is presented as a new process monitoring tool with the ability to provide information not accessible through traditional SLS process monitoring methods.Principles of Optical Coherence TomographyOptical coherence tomography (OCT) is an established medical imaging technique basedon low coherence interferometry commonly used in biomedical optics and medicine. OCT is used in this field and others for its ability to provide subsurface visualization of translucent or turbid materials. OCT produces cross-sectional and volumetric images from magnitude and echo time delay data of backscattered radiation.

OCT produces cross-sectional and volumetric images from magnitude and echo time delay data of backscattered radiation. Samples can be imaged in real time with typical axial resolutions of 1-15 microns.1397A single OCT scan at a point on the surface of the sample, is known as an A-scan andyields depth information at this point. A collection of subsequent A-scans is called a B-scan and yields cross sectional information for the sample. Subsequent B-scans can be combined to yield a 3D representation of the sample as shown in Figure 1 [3].Figure 1 (Left) A single A-scan is shown. (Middle) A cross sectional B-scan is shown with intensity data shown in false color. (Right) Many B-scans have been combined to form a volumetric rendering of the sample. [3]Figure 2 illustrates some basic OCT concepts. In OCT, light from a laser is split into“sample” and “reference” beams.

[3]Figure 2 illustrates some basic OCT concepts. In OCT, light from a laser is split into“sample” and “reference” beams. The reference beam is split from the source light and serves as a reference time delay for light returning from the sample. The sample beam is forward scattered and possibly absorbed as it propagates into the media of interest. Some radiation entering the sample is backscattered and re-combines with the reference beam forming an interference between these beams that is measured by a photodetector. From this measurement, spatial information about the sample can be obtained. Interfaces where the index of refraction changes suddenly (large longitudinal gradient), including variations in material phase and material type are often detectable with OCT.There are several different types of OCT systems. An optical fiber-based swept sourceOCT system (SS-OCT) was used for the experiments in this paper. Detailed information about other types of OCT can be found in [3].

An optical fiber-based swept sourceOCT system (SS-OCT) was used for the experiments in this paper. Detailed information about other types of OCT can be found in [3]. Additional SS-OCT specific background information is provided below.In SS-OCT systems, the wavelength of the radiation emitted by the laser changes in timeas it sweeps across a frequency range during each A-scan. The light in the sample arm is configured to travel slightly longer than that in the reference arm. The difference in pathlength between light in reference and sample paths causes the phase of the interference fringes to vary as the optical source frequency (ν) varies. The phase (φ) of the interference between reference and sample paths varies as φ =2πcτν where c is the speed of light ν is the optical frequency and τ is the time delay between sample and reference paths.

Because light in the sample beam travels variable path length delays (τ) depending on how far into the sample it penetrates before back1398reflecting, a Fourier transformation is necessary to allow the individual path length delays in the interference signal to be separated out yielding the axial scan information (echo magnitude vs. time delay).KeySSSwept Source LaserBSBeam SplitterREFReference ArmSMPSamplePDPhotodetectorDSPDigital Signal ProcessingFigure 2: A schematic of a simple two-beam swept source OCT system (SS-OCT) [4]Stifter reviews several applications of OCT which including inspection of polymericinjection molded parts of polyamide and PEEK materials [5]. Quite recently, OCT was used for ex-situ analysis of SLS manufactured parts by Guan et.al. [6]; however, to the authors’ knowledge OCT has not been used as a process monitoring tool during an SLS build.

[6]; however, to the authors’ knowledge OCT has not been used as a process monitoring tool during an SLS build. In this paper, OCT is investigated as a tool for in-situ process monitoring during the selective laser sintering of nylon 12. To this end, ex-situ analysis was first performed to verify the feasibility of OCT for in-situ process monitoring.Ex-Situ Part Examination: Experimental SetupFor the ex-situ part imaging, the OCT system used a commercially available 30 mWAXSUN swept source laser with a center wavelength of 1310nm and a wavelength tuning range of 125nm at a 100 kHz sweep rate. The axial scan resolution is approximately 3.75 μm/pixel in air and the lateral resolution is approximately 16 μm/pixel. The maximum scan depth is approximately 4 mm in air. Each B-scan is made up of 512 A-scans. All of the parts imaged for the ex-situ part examination were built in UT Austin’s Laser Additive Manufacturing Pilot System (LAMPS) machine.

Each B-scan is made up of 512 A-scans. All of the parts imaged for the ex-situ part examination were built in UT Austin’s Laser Additive Manufacturing Pilot System (LAMPS) machine. The LAMPS machine was developed at UT Austin for the study of SLS process control in high melting-point temperature polymers [9]. A more detailed description of the LAMPS machine is provided later in this paper.Ex-Situ Part Examination: ResultsThe following selectively laser sintered parts were imaged as part of the preliminary exsitu OCT analysis: a nylon 12 tensile bar specimen with a ledge, half of a nylon tensile bar specimen in which layering was visible, and a PEEK part with noticeable surface pitting.1399Nylon 12 Tensile Bar w/ LedgeFirst, a nylon 12 part with a ledge was imaged. The ledge was created inadvertently asthe SLS powder roller moved the previously sintered layers slightly while spreading subsequent powder layers.

The ledge was created inadvertently asthe SLS powder roller moved the previously sintered layers slightly while spreading subsequent powder layers. This part was selected because the ledge allows one to ascertain if the OCT system is able to penetrate through a few layers of the material before the backscattering of the laser is indistinguishable from noise. Figure 3 shows visual images of the described part.Figure 3: (Left) View of tensile bar from above (Right) The ledge is apparent on the left of the image when looking at the tensile bar from the end.Figure 4 shows the B-scan cross section of the tip of the ledge. The bottom boundary ofthe ledge is clear up to at least the red dotted line in Figure 5, which is approximately 0.36 mm or 3.6 layers in this particular part.

The bottom boundary ofthe ledge is clear up to at least the red dotted line in Figure 5, which is approximately 0.36 mm or 3.6 layers in this particular part. Because this was merely a feasibility study in preparation for the in-situ analysis, and did not include surrounding un-melted powder, precise criteria were not used to identify the maximum depth of resolution of this ledge. The surface depth is assumed to be indicated by the local maximum of intensity values versus depth in each A-scan. In absence of a measurement, the index of refraction of the sintered material was assumed to be 1.53 and 1.00 for the air [7].Figure 4: (Left) A visual image is provided for reference with the approximate region of the OCT scan indicated by the square.

(Right) A single B-scan of the tensile bar ledge is shown.1400Bottom SurfaceTop SurfaceFigure 5: A scaled intensity value from a single A-scan is shown with the top and bottom surfaces sintered parts are indicated.Half of Nylon 12 Tensile BarA half of a nylon 12 tensile bar specimen was also imaged. Layering was visible by eye on the outside of the part. Layering was also visible using OCT. Figure 6 shows a visual image of the part described. Figure 7 shows the OCT image of the layered half tensile bar and Figure 8 shows the average intensity values for a 20 A-scan region.Layering VisibleBuild Layer DirectionImaged SurfaceFigure 6: A Nylon 12, Layered Half Tensile Bar Specimen. The surface imaged with OCT was the vertical surface closest to the viewer.1401Figure 7: A single B-scan of the tensile bar is shown. Horizontal layers of bright intensity are visible just beneath the surface.

Horizontal layers of bright intensity are visible just beneath the surface. The average intensities of the A-scans between the dotted lines are given in Figure 8.Top SurfaceStriations VisibleFigure 8: The average intensity of 20 consecutive A-scans is shown. The horizontal layering is evident after the surface peak. The averaging was performed to reduce the effect of noise.1402In Figure 8, 6 somewhat regularly spaced striation peaks are present. It is thought that these striations correspond to the individual build layers, but additional analysis is required in order to confirm this.PEEK Part with Noticeable Surface PittingA part made from polyether ether ketone (PEEK) with noticeable pitting on the surfacewas imaged next. Figure 9 shows a visual and OCT image of the PEEK part.1mmFigure 9: (Left) A visual image of a previously broken tensile bar made of PEEK powder with poor surface quality. (Right) Pits can be seen in surface quality.

(Right) Pits can be seen in surface quality. Potentially pits down to the size of 16 microns (the width represented by a single pixel) can be seen with the current OCT configuration.In-Situ OCT: Experimental SetupLAMPS MachineThe in-situ imaging of polymer powder and sintered material was done by installing anOCT system on the LAMPS machine. The LAMPS (Laser Additive Manufacturing Pilot System) machine was designed and built at The University of Texas at Austin as an open architecture SLS printer to allow increased monitoring and control of the process variables. The machine is outfitted with visual and infrared cameras and the ability to add and remove additional data acquisition elements (such as OCT) as needed. A more complete description of the LAMPS System can be found in Wroe et al. [8] and Fish et al. [9]. The right side of Figure 10 has been provided from Wroe et al.

A more complete description of the LAMPS System can be found in Wroe et al. [8] and Fish et al. [9]. The right side of Figure 10 has been provided from Wroe et al. showing the LAMPS machine [8].1403Figure 10 – (Left) External View of LAMPS CAD geometry (Right) LAMPS Build Chamber Cross SectionOCT SetupFor the in-situ OCT imaging, a 20 mW Santec HSL-2000 swept source laser with acentral wavelength of 1310 nm and a sweep range of 60 nm at a scan rate of 20 kHz was used. The axial scan resolution is approximately 11 μm/pixel in air and the lateral resolution, with a necessarily long focal length lens, is approximately 186 μm/pixel. The optics of the LAMPS system were modified such that the OCT laser was boresighted with the CO2 laser as shown in Figure 11.This setup aligns the OCT beam to scan wherever the laser is pointing at all times.

Inactual implementation for this paper, the OCT beam scanned about 1 mm behind where the CO2 laser was melting the surface powder.The sample arm pathlength changes in relation to the reference arm pathlength based onwhere the CO2 laser is pointing on the build surface. This causes the flat powder surface to appear curved in the OCT images. The longer path lengths farther from the center of the build area appear lower in the figures below. The OCT signal also weakens near the edges of the build area as the powder moves farther from the plane of best focus, however, both issues could be mitigated by using a field-flattening lens in the sample path.

The present analysis was limited to data collection over a small region near the center of the build area so signal roll-off is not a major concern, but the flat powder surface does appear slightly curved in Figure 13 and Figure 15.This boresighted OCT setup produced OCT images that were noisier than those produced from the ex-situ part analysis. This is because the long focal length lens in the boresighted OCT setup leads to a reduction in the solid angle of back reflected radiation coupled back into the system. In addition, the large spot size combined with the surface variation of the sample produced a reduced signal amplitude. Using a more powerful laser in the boresighted system would help to produce higher quality images in the boresighted setup in the future.1404Figure 11 - Schematic of the Boresighted OCT OpticsIn Situ OCT: ResultsThe in situ tests were run in the LAMPS machine using unrecycled polyamide 12 powder(PA 650, ALM).

A single layer test designed to look at how a single scan line changes in time was run as well as a multilayer test looking at a few layers of sintered material.Single Layer Scan LineFor this test, two markers were made approximately 4 cm apart by lasing a single pointwith 2.5 W for 0.1 seconds. This caused an easily recognizable divot in the surface of the powder. The marking holes were used to help identify the location of the OCT beam during post processing. Then, a pre-lase OCT scan was taken followed by the lasing of a 2.5 cm section in the middle of the two marking holes with the CO2 laser. After lasing, the OCT beam was scanned back and forth across the freshly lased region and marking holes 100 times in about 3.2 seconds. This is shown in Figure 12. This test was run several times using different lasing powers.1405Figure 12: Diagram of the build surface from above for the single layer scan line test. First, two marking holes were made.

This test was run several times using different lasing powers.1405Figure 12: Diagram of the build surface from above for the single layer scan line test. First, two marking holes were made. Then, a single line was lased in between the two marking holes (blue area). Afterwards the OCT beam scanned back and forth (red arrow) over the marking holes and lased area 100 times in about 3.2 seconds.The results from a high laser power, 32.5 W, are shown here. Although this is more thantwice the laser power generally used during builds, the high laser power was able to accentuate some phenomena that are present, though less noticeable, at lower laser powers.Figure 13 shows the pre and post-lasing OCT scans. Both images have had a non-localmeans de-noising filter applied [10]–[12]. The post-lase OCT image shows a melted region where the laser melted the powder together. The melted region itself is mostly transparent in the image.

The post-lase OCT image shows a melted region where the laser melted the powder together. The melted region itself is mostly transparent in the image. This makes sense because there are no interfaces in a pure liquid for radiation to reflect back from. However, the top liquid surface is clearly visible. The melted region is deeper on the right hand side. It was already known from previous work that a hot spot was created at the beginning of each scan line due to uneven heating while the laser accelerated. With OCT, the additional depth of the meltpool caused by overheating can be measured. This deeper melted region from the hot spot could be seen on a reduced scale at laser powers more typical of actual build conditions (not shown). The ghost line is an undesired effect of a piece of dust or other contaminant in the OCT system setup, causing a reflection which is not representative of the sample being imaged.

The ghost line is an undesired effect of a piece of dust or other contaminant in the OCT system setup, causing a reflection which is not representative of the sample being imaged. This effect is local and is ignored in the present analysis.Although the lower laser powers didn’t produce a melt region as transparent as the one in Figure 13, they did show a region of reduced intensity. This can be seen in the multilayer test as well (see Figure 15).1406Pre-Lase OCT ScanPost-Lase OCT ScanFigure 13: (Above) A single B-Scan of the test area before lasing with the CO2 laser. (Below) A B-scan of the test area 3.2 seconds after lasing with the CO2 laser.Multilayer Scan PatternFor the multilayer scan test, several layers of a 1” by 1” square were lased. In each layer, several marking holes were created, then a pre-lase OCT scan was performed followed by lasing of the square area, followed by a post-lase OCT scan.

In each layer, several marking holes were created, then a pre-lase OCT scan was performed followed by lasing of the square area, followed by a post-lase OCT scan. Between layers, a new layer of powder was dropped, heated to approximately 140oC and spread over the build surface. Figure 14 shows a diagram of a post-lase OCT scan for the multilayer test.1407Figure 14 - Diagram of the build surface from above for a post-lase OCT scan of a single layer of the multilayer test. The marking holes and lased area are shown in blue. The first 3 OCT B- Scans are shown in red. 50 B-Scans make up each pre and post-lase OCT scan.Figure 15 below shows corresponding B-scans on pre and post-lase OCT scans on thefirst layer as well as the post-lase OCT scan on the 2nd layer. The region that was lased lies in between the jump and mark delay regions. The jump and mark delays are brief pauses in laser movement after either a jump or a mark period respectively.

The region that was lased lies in between the jump and mark delay regions. The jump and mark delays are brief pauses in laser movement after either a jump or a mark period respectively. Their original purpose is to ensure that the laser fires only on the intended area to be sintered. They produce readily recognizable regions where subsequent A-scans are relatively constant because the sensor is not moving and data is being duplicated in time.In the 1st layer post-lase OCT scan (middle image), curling can be seen at the left of thelased region. The sintered layer also shows up as a reduced intensity layer on top of the un- sintered powder. In the 2nd layer post-lase OCT scan, the top 2 sintered layers are visible, although the boundary between the lower sintered layer and the un-sintered powder is not as defined as the boundary between the two sintered layers. The build failed after the 2nd layer due to part curl.

The build failed after the 2nd layer due to part curl. The build was repeated at a reduced laser power and similar results were seen, but with layer interfaces being slightly less defined. Layers below the top 3 were not readily differentiated by visual inspection. This could likely be improved with a more powerful OCT laser.Future WorkThe fact that layers can be differentiated from each other and from un-sintered powdersuggests that one could quantify the level of inter-layer bonding and perhaps predict where a part will fail. This will be analyzed in future work. Additionally, the fact that curling can be detected before the curl becomes severe enough to cause the build to fail indicates that it may be possible to correct a curl by using a lower laser power on the curling area or by adjusting the heating elements of the build chamber to reduce future curl.

The LAMPS machine is ideally set up for both of these types of experiments and these will be examined in the future.1408Figure 15 – Three corresponding B-Scans are shown. The individual layers as well as part curl are evident in the OCT images. (Top) 1st Layer Pre-Lase OCT Scan (Middle) 1st Layer Post-Lase OCT Scan (Bottom) 2nd Layer Post-Lase OCT Scan1409ConclusionThe feasibility of OCT as a process monitoring tool was confirmed through ex situanalysis. Specifically, striations and surface roughness were observed in ex-situ part analysis. Then, for the first time an OCT system was bore sighted with the CO2 laser on the LAMPS SLS machine and OCT data was taken of the powder and sintered material in-situ. It was observed that part curl can be detected before it causes the build to fail.

It was observed that part curl can be detected before it causes the build to fail. Additionally, the top layers of the sintered material were resolved and differentiated from un-sintered powder.AcknowledgementsThe authors would like to acknowledge the instrumental research support of General Dynamics Information Technology, Inc. (GDIT) PO# 08ESM753983: Laser Additive Manufacturing - Pilot Scale System (LAMPS) (PI: Scott Fish); the Office of Naval Research (ONR) Award# N00014- 12-1-0811/N00014-16-1-2393: Physics and Cyber-Enabled Manufacturing Process Control for Direct Digital Manufacture of Metal Parts (PI: Joseph Beaman); and the National Science Foundation (NSF) Award # CNS-1239343: CPS: Synergy: Cyber Enabled Manufacturing Systems for Small Lot Manufacture (PI: Joseph Beaman) for their support of this research.References[1] T. Wohlers, T. Caffrey, and R. I. Campbell, Wohlers Report 2016: 3d Printing and Additive Manufacturing State of the Industry, 1st ed.

...SITU PROPERTY MEASUREMENTS ON LASER...DRAWN STRANDS OF SL 5170 EPOXY AND SL 5149 ACRYLATET. R. Guess and R. S. Chambers Sandia National Laboratories Albuquerque, NM 87185-0958ABSTRACTMaterial behavior plays a significant role in the mechanics leading to internal stresses and, potentially, to distortion (curling) of parts as they are built by stereolithography processes that utilize photocuring resins. A study is underway to generate material properties that can be used to develop phenomenological material models of epoxy and acrylate resins. Strand tests are performed in situ in a 3D System's SLA-250 machine; strands are drawn by either single or multiple exposures ofthe resin to a laser beam. Linear shrinkage, cross-sectional areas, cure shrinkage in cantilever beam forces and stress-strain data are presented.

Linear shrinkage, cross-sectional areas, cure shrinkage in cantilever beam forces and stress-strain data are presented. Also, specimens, built with different draw patterns, are compared.the curlINTRODUCTIONDuring the stereolithography build process, parts are built layer by layer using a laser to selectively cure slices of photocuring resin. As the laser spot passes over the surface of the resin, the ensuing chemical reaction causes the resin to shrink and stiffen during solidification. When laser paths cross or when new layers are cured on top of existing layers, residual stresses are generated as the cure shrinkage of freshly gelled resin is constrained by the adjoining solid material.

These internal stresses can cause significant out-of-plane curling in the compliant material.A program is underway at Sandia National Laboratories to develop and to evaluate the use of engineering tools for modeling curl and for studying parameters that might improve overall build accuracy in stereolithography parts [1]. These tools include phenomenological material models of the solidifying resins and a 3-D finite element model. that incorporates time varying material behavior, laser path dependence, and structural linkage [2]. Methods developed for in situ response of polymer strands drawn during single (or multiple) passes of a He-Cd laser in 3D system's SLA-250 stereolithography apparatus have been reported along with initial experimental results [1]. Additional strand data are presented below for two Ciba Giegy photocurable resins- Cibatool® SL 5170 epoxy and Cibatool® SL 5149 acrylate. The strand specimens were drawn in an SLA-250.

Additional strand data are presented below for two Ciba Giegy photocurable resins- Cibatool® SL 5170 epoxy and Cibatool® SL 5149 acrylate. The strand specimens were drawn in an SLA-250. The data include linear shrinkage, cross-sectional areas, cure shrinkage forces and initial modulus data of strands following single and multiple laser hits. In addition, out-of-plane curl data for acrylate cantilever beam specimens built with different draw patterns are discussed.TEST PROCEDURES AND RESULTSLINEAR SHRINKAGE- Time history linear shrinkage of SL 5170 epoxy and SL 5149 acrylate strands were obtained by using a non-contact video method to measure the change in length of individual strands after gelation following a single laser hit. These data were initially reported in [1]; representative curves for one inch long specimens are shown in Figure 1. Note134that for a single laser exposure, the epoxy has a linear shrinkage of about 1.5% and the acrylate approximately 1.0%.

Note134that for a single laser exposure, the epoxy has a linear shrinkage of about 1.5% and the acrylate approximately 1.0%. Shrinkage is complete about three minutes after gelation.AREA MEASUREMENTS- The metallography technique developed to determine the shapes and areas of SL 5170 epoxy and SL 5149 acrylate strands was described in [1]. The same technique was used to measure epoxy. and acrylate strand cross sections following single and multiple laser hits; typicalbinaryimagesare shown in Figure The strands were not recoated with fresh resin between hits inthe multiple hit experiments. Curves of strand as a function of number of laser hits are plotted in ·Figure 3.

Curves of strand as a function of number of laser hits are plotted in ·Figure 3. Note that the cross sectional area of the epoxy increases at a higher rate than the acrylate area does.FORCE MEASUREMENTS- Fixtures developed for in situ measurement of forces during cure shrinkage, step displacementloading, and stress relaxation tests of strands are schematically illustrated in Figure 4; more detail is given in [1]. A drawn strand attaches to tapered tabs that are just below the resin surface; one tab is connected to an extension from a miniature force gage and the other tab is connected to a platform whose movement can be controlled by an electronically driven micrometer. A 1.0 inch strand was used in all force tests. The 150 gram load force gage, calibrated to 20 grams full scale has a resolution of ±0.02 gram. An LVDT (± 0.05 inch range) is used to monitor strand displacement imposed by the micrometer whose displacement rate and amplitude can be controlled.

An LVDT (± 0.05 inch range) is used to monitor strand displacement imposed by the micrometer whose displacement rate and amplitude can be controlled. In all 0.5% step strain tests, strands are strained at a rate of about 0.01 S-l.situ force tests are performed to measure (1) cure shrinkage forces developed after single or multiple laser hits on single strands, (2) stress-strain response when a 0.5% step strain is imposed on a strand at different elapsed times after the strand is drawn, and (3) force-time relaxation response after 0.5% step strain. Initial apparent modulus values are based on the initial slope of force-strain curves and the cross sectional area of the strand. These values of modulus are integrated averages over the strand cross section, whose properties are nonhomogeneous due to a cure gradient resulting from attenuation of laser energy the resin.

These values of modulus are integrated averages over the strand cross section, whose properties are nonhomogeneous due to a cure gradient resulting from attenuation of laser energy the resin. The cure gradient determines the extent of reaction (degree of cure) at different locations in the strand cross section.SL 5149 acrylate data- A typical force versus time response of an SL 5149 acrylate strand which was hit with 4 laser exposures prior to applying 0.50% step strain is shown in Figure 5. This curve illustrates that no shrinkage forces are detected after the 1st laser hit, but that cure shrinkage forces are measured following additional linear shrinkage associated with multiple hits can not be measured because the strands are attached at both ends and cannot shorten. The stress-strain responses during the 0.5% step strain provide initial modulus values.laser hits. The actualexposures. number of laser hits increases.

The stress-strain responses during the 0.5% step strain provide initial modulus values.laser hits. The actualexposures. number of laser hits increases. A small region of Figure 6 is expanded in Figure 7 to show in greater detail the shape of the force-time response during and following a laser hit. It is of interest to note that immediately following a laser hit, the force initially decreases, possibly due to thermal expansion of the strand from laser heating. The strand then undergoes additional cure shrinkage producing a force that initially increases rapidly before leveling off.Figure 6 clearly shows that the acrylate continues to shrink for at least 20 magnitude of each additional cure shrinkage force increment decreases as135modulus ofstrands as a function of theof laser8.modulus calculations account for the strand area associated withinitial modulus increases from about 1000 psi for ato near 110,000a acrylate c"'r<:lnrlcexposed to 28 laser hits.

It was reported following a single exposure to application of modulus (degree is not recoated with[1] that the initial modulus with increasing time interval step strain. Acrylate strands continue to shrink and to lnf'1I"Pt;H~Pdid notcure) following multiple laser hits, for at least 28 hits.between laser exposures.~~~~~~~ .. No shrinkage forces were detected for SL 5170in testsin which specimens were subjected to as many as 8 laser hits. However, forces were .11.. when strands were subjected to a 0.5% step strain. Strands were strained at ................""'... '''''' times following the laser draw. Results times up to 240 minutes show that both initial stiffness and peak force increase with elapsed times appear to be influenced by strain.

'''''' times following the laser draw. Results times up to 240 minutes show that both initial stiffness and peak force increase with elapsed times appear to be influenced by strain. Typical stress..strain responses for multiple hit strands are plotted in strain curves for 1, Strands with more hits curves are compared with data seem to illustrate that the initial modulus not the number,""'tests on strands of a singlehitHowever,tests on strands after multiple laser hits, initial modulus does not number of hits but by the elapsed time from the 1st step stress..toand 3 laser hits are about identical (all with an elapsedlonger elapsed times had stiffer responses. Modulisingle strike data as a function5170 epoxy is dependent on '-'........11-1..:>""""IJ""' .....' .....hits as is the SL 5149 acrylate.of 11 were built and the resulting out..of..plane curl was measured...

SL 5149 acrylate cantilever beamshown in laser beam vector paths were used to build the specimens. laser was always path recoat was cantilever beam is 6 layers thick. layer receives (either 1 or 2). There was no resin recoat between the 1st and 2nd hits of those artifacts which had aone build style,the 1..(length..wise) direction of thethevectors for each layer; the center section of(transverse) directioncantilever. is 8thick andother variable is the number of laser hitson each layer. of the individual layers and betweenall cases, there was a three minutehit on on eachsamebeams of each build style (all with 1of the four were postcured in a UV oven for 30 .............1............... ..:>. photographed using an optical microscope. Its was used to measure·the out..of..plane resultsdrawn with l ..direction and photographs show~Vj::,)L""'I-.1J...." clearly show the(a) results from.....""'·n " ..... draw ....

Its was used to measure·the out..of..plane resultsdrawn with l ..direction and photographs show~Vj::,)L""'I-.1J...." clearly show the(a) results from.....""'·n " ..... draw .... <:1 ...1·0.........12(c) show that the bottom edge curl is about 0.016 specimens. Neither the draw style nor the postbeam for or distribution of curl in the acrylate beams.at the tip of the 0.15 step influenced the ............,f",......Jl ..........."'"136The fact that the draw direction did not affect curl in SL 5149 acrylate is probably because the draw time of a single layer is on the order of a few seconds, a very short time relative to the minutes necessary for full shrinkage to occur (see Figure 1). Alternate draw styles and time delays that may cause cantilever beams to have different amounts of curl are planned for future tests.

Alternate draw styles and time delays that may cause cantilever beams to have different amounts of curl are planned for future tests. If those results are successful, cantilever specimens will also be built with the SL 5170 epoxy.SUMMARYIn situ property measurements were made on single strands of SL 5170 epoxy and SL 5149 a single laser exposure, epoxy strands shrank approximately acrylate photocurable resins. 1.5% and the acrylate strands about 1.0%. The cross sectional area of strands increased with additional laser hits (no recoatingwith liquid resin between hits). The increase in area with laser hits was greater for the epoxy than for the acrylate. No cure shrinkage forces were measured after the first laser hit in either epoxy or acrylate, and none was measured after multiple laser hits in the epoxy. This suggests that the epoxy either does not continue to shrink after the 1st hit or that shrinkage forces cannot overcome strand stiffness.

This suggests that the epoxy either does not continue to shrink after the 1st hit or that shrinkage forces cannot overcome strand stiffness. On the other hand, the acrylate develops additional cure shrinkage forces for 20 or more laser hits. In force· relaxation tests, a strand is drawn and then a 0.5% step strain is applied after different number of laser hits and/or elapsed time between the first hit and the step strain. The acrylate initial modulus is a function of the number of laser hits, but not the elapsed time. In contrast, epoxy modulus is a function of elapsed time, but not the number of laser hits.Procedures were developed to measure out-of-plane distortion (curl) of cantilever beam specimens. Beams (depth of 0.060 inch and length of 0.15 inch) drawn with two different build styles had curl of about 0.016 inch at the tip. Neither the draw style nor a post cure step influenced the magnitude or distribution of curl in the acrylate beams.

Neither the draw style nor a post cure step influenced the magnitude or distribution of curl in the acrylate beams. The draw time was so short that essentially no shrinkage occured during the drawing of a single layer by either of the two draw patterns used.REFERENCES D. Hinnerichs, G. D. McCarty, and R. N. Shagam T. D., "Epoxy and Acrylate Stereolithography Resins: In-Situ Measurements of Cure Shrinkage and Stress Relaxation," Proceedings of the Sixth International Conference on Rapid Prototyping, June. 4-7,1995, Dayton, Ohio.[1]. T. R. Guess, R. S. Chambers,[2]. Chambers, R. S., Guess,D., "A Phenomenological Finite Element Model of Part Building in the Stereolithography Process," Proceedings of the Sixth International Conference on Rapid Prototyping, June. 4-7, 1995, Dayton, Ohio.R., and Hinnerichs,ACKNOWLEDGMENTSThis work was performed at Sandia National Laboratories and supported by the U.S.Department of Energy under contract DE-AC04-94AL85000.

Brian Pardo, Daryl Reckaway and operated the SLA-250 equipment, drew the strands and performed the force Mark relaxation tests. Dale McGuffin prepared the metallographic samples and made the area and curl measurements.1372.0".-..~o '-'" ~ 1.5 ~ z 0::: 1.0 I C/)0:::I1i 0.5 zI0.0~~~'"NrHi~ ...EP pXY ,.-i••~ - ~ ~-n ~--.'" ..~- Jj"""....ACRYLATE......./r..J',......0.010.11101001000TIME (5)Figure 1. Linear shrinkage of SL 5170 epoxy and SL 5149 acrylate strands following a singlelaser exposure.SL 5149 ACRYLATESL 5170 EPOXYA=.000160 in. 2 W=.0102 in. D=.0196 in.A=.000142 in. 2 W=.0117 in. D=.0167 in.A=.000203 in.2 W=.OIII in. D=.0264 in.A=.000284 in.2 W=.0163 in. D=.0293 in.Figure 2. Cross sections of SL 5149 acrylate and SL 5170 epoxy strands for different number oflaser hits. Arrow indicates direction of laser hit. Strands were not recoated between hits.

~ .005..-2-DirectionJJJ- ..o .000 ....................... ::> 0.00+-..a.....IL....I.......- I -..............................0.050.100.15DISTANCE ALONG BEAM (in)(c)Figure 12. Out-of plane curl of six -layer acrylate cantilever beams (about 0.010 inch thick layers) for two draw patterns and cure conditions: (a) side view of a beam built with a 2-direction (transverse) draw pattern, (b) a beam built with a I-direction pattern, and (c) comparisons of bottom surface curl versus distance along the beam for specimens of the two build patterns, without and with postcure.141142.

lid Freeform Fabrication 2016: Proceedings of the 27th Annual International Solid Freeform Fabrication 2016: Proceedings of the 26th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIN-SITU QUALITY INSPECTION OF LASER POWDER-BED FUSION USING HIGH- RESOLUTION VISUAL CAMERA IMAGESMasoumeh Aminzadeh, Thomas KurfessWoodruff School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332AbstractIssues of part quality in terms of quality of fusion and formed porosity are widely known and stated as some of the important challenges with laser powder-bed fusion (LPBF) process. This paper addresses the in-situ inspection of layer-wise part quality using visual camera images. High- resolution visual images are captured from each layer of the part during LPBF process.

This paper addresses the in-situ inspection of layer-wise part quality using visual camera images. High- resolution visual images are captured from each layer of the part during LPBF process. The imaging and illumination setups are developed such that the produced images visualize detailed surface characteristics of each layer of the build such as fused seams, as well as the individual formed pores. To enable automated inspection of these images, appropriate image processing algorithms are developed to detect individual pores formed in each layer. In addition to detection of individual pores, intelligent pattern matching algorithms are developed, trained, and implemented to identify porous regions from non-porous layers. The surface characteristics of the layers as visualized in camera images can also provide a measure of quality of fusion and the energy of the layer, and an estimated level of porosity.

The surface characteristics of the layers as visualized in camera images can also provide a measure of quality of fusion and the energy of the layer, and an estimated level of porosity. Discussion on characterization of the surface quality in terms of roughness, quality of fusion, and the energy of the build will be made. The results of the automated image analyses provide useful feedback for in-situ process modification as well as part quality assessment.IntroductionThe need for developing techniques and approaches to improve and control the quality of parts made by metal powder-bed additive manufacturing (AM) is widely known and stated. The high chances of formation of pores in metal AM techniques such as laser powder-bed fusion (L- PBF) can lead to fabrication of parts that do not meet the requirements for quality and structural properties.

Therefore, to assure the quality of parts, it is required to develop strategies to monitor and control formation of defects and pores in the part, especially in an in-situ fashion during the build.The majority of the past work on quality control in metal AM focuses on real-time process control as a means to control the physical phenomena associated with powder fusion process and consequently prevent formation of pores. However, despite partial improvements achieved by real- time process control, due to the high level of complexity of powder fusion process and its governing factors, defects and pores still form in the parts. Therefore, an alternative or complementary approach is sought which is based on directly monitoring or inspection of the defects such as pores.

Therefore, an alternative or complementary approach is sought which is based on directly monitoring or inspection of the defects such as pores. The results of defect inspection can then help with making decisions regarding the acceptance of the built portion of the part, process control, and taking corrective actions to fix the defects if possible.1412While measurements such as 2D thermal data or acoustic waves might provide insight into the quality of the part, 2D visual camera images have the capability to visualize defects as well as individual pores provided appropriate illumination arrangements and setup. There is a very limited number of work on in-situ inspection in metal AM, or specifically powder-bed AM, using visual camera images. Craeghs et all., [1], used a visual camera to build a powder deposition monitoring system in L-PBF. They detected the effect of a worn or damaged coater that can be regarded as specific type of process error that can lead to part defects.

They detected the effect of a worn or damaged coater that can be regarded as specific type of process error that can lead to part defects. Kleszczynski et al., [2], used a 29 megapixel visual camera mounted outside the machine to visually examine process errors in EBM. They studied the effect of factors such as input energy on formation of elevations at edges which they manually observed in high-resolution camera images. In another subsequent work, [3, 4], they used their imaging system to capture images with 24 μm/pixel on-part resolution to inspect elevations in the part.

In another subsequent work, [3, 4], they used their imaging system to capture images with 24 μm/pixel on-part resolution to inspect elevations in the part. They captured camera images from the deposited powder layer and used constant value thresholding to detect elevated regions in the part.Aminzadeh et al., [5, 6], proposed a framework based on machine vision for inspection of geometric errors as well as pores; they designed and implemented image processing algorithms to automatically detect the geometric errors and inspect the dimensional accuracy of the part from high-resolution camera images. In a subsequent work, [7], they used a 9 megapixel visual camera mounted inside the build chamber perpendicular to the build platform that provided images with sufficient contrast between the powder and fused regions at a very detailed 7 μm/pixel on-part resolution.

They developed automatic image segmentation algorithms that detected the geometric cross section of the parts with 80 μm detection error which was smaller than the laser scan diameter.Foster et al. [8], used a 29 megapixel camera mounted inside the machine at an angle from the build platform. They were able to capture images with 50 μm/pixel resolution and 15 μm/pixel resolution depending on the utilized camera lens. The fused regions are visualized with relatively low contrast in the captured images. However, despite the high resolution, the formed defects and detailed surface characteristics were not yet visible in the images. They studied the effect of illumination on image characteristics to provide images that can later be used for in-situ inspection purposes.This paper describes the system and the work on development of a new imaging setup with the collaboration of the Edison Welding Institute (EWI) at EWI.

Unlike the previous and existing work on development of imaging setup for powder-bed AM, this system resolves, to a significant extent, the challenges associated with the illumination and captures images that visualize detailed surface features such as individual pores as they appear in microscopic images. Appropriate image processing algorithms are then developed to detect pores from the part surface automatically. In addition to detection of individual pores, an intelligent classification algorithm is developed to identify porous versus non-porous layers.1413Imaging Setup and Discussion on Images CharacteristicsFigure 1 shows a close-up of the build chamber of a selective laser melting (SLM) machine that was built and customized at EWI. An IPG 600 watt single mode fiber laser was used as the lase source. As seen in the Figure, in this customized machine, the laser source is mounted with some shift to the left at the top of the build platform.

As seen in the Figure, in this customized machine, the laser source is mounted with some shift to the left at the top of the build platform. This feature provides room right at the top of the build platform for mounting sensors.An 8.8 megapixel USB Digital Camera with high focus lenses was mounted directly at the top of the build platform looking down perpendicular to the build. The camera resolution and the captured image size is 4096×2160 pixels. The following features of this imaging setup are noteworthy. The camera is mounted inside the machine right at the top of the build platform. This significantly reduces the distance between the camera and the build platform. Utilizing lenses with a small focal length would allow capturing images with very high on-part resolution in order of several micron pixel size. Additionally, the camera captures pictures at a perpendicular angle from the build platform which results in no perspective effects.

Additionally, the camera captures pictures at a perpendicular angle from the build platform which results in no perspective effects. Consequent, the picture is uniformly scaled throughout the image with the same pixel size all over the image. Therefore, there would be no need for camera calibration to fix for shape and size distortions due to perspective effects. Camera calibration would have added more computational complexity and would introduce new errors and uncertainties due to machine vibration [9].Figure 1- Close up of the build chamber of the machine and its components.The other important component of the imaging setup in addition to the camera is the illumination (light) source. The location and orientation of the camera and the light source play critical role in the quality of imaging. The best orientation for the camera is perpendicular to the build to remove the perspective effects and the need for camera calibration that was met in this imaging setup.

The best orientation for the camera is perpendicular to the build to remove the perspective effects and the need for camera calibration that was met in this imaging setup. The camera was mounted conveniently on a stationary fixture connected to the wall of the chamber. It was mounted such that it doesn’t block the laser path. It was later examined that the distance of the camera from the build in this setup is not optimal and better image might be captured if the camera be put somewhat closer to the build depending on the lens focal length.1414However, this current position for camera was mainly determined based on the physical limitations.To mount the light source, initially a single light source in form of an adjustable, white, high intensity LED spot light, was mounted at the top of the build platform emanating at an angle, as shown in Figure 1. The level of brightness was adjusted until rather clear images of the powder bed could be captured.

The level of brightness was adjusted until rather clear images of the powder bed could be captured. To examine image quality, builds with square cross sections were fabricated for several layers. Images were captured and acquired at leach layer. Figure 2 shows an example of images captured with this illumination setup.With an 8.8 megapixel camera with 4096×2160 pixel image sizes, the on-part pixel size for images was measured as 7 µm. Before using the 8.8 megapixel camera in this imaging setup at EWI, a 2 megapixel camera with 1080×1920 pixel image sizes was utilized that led to the on-part pixel size as 20 µm with the same camera mounting location. This is whereas the resolution achieved by Kleszczynski et al. [3, 9] with a 29 megapixel camera with image size of 4234×4234 pixels was 24 µm/pixel, and the resolution achieved by Foster et al. [8], using a 29 megapixel camera, reached 50 or 15 µm/pixel, depending on the utilized lens.

[8], using a 29 megapixel camera, reached 50 or 15 µm/pixel, depending on the utilized lens. This better resolution achieved in the imaging setup described in this work, 7 µm/pixel using an 8.8 Megapixel camera, is due to the smaller distance of the camera from the build and the perpendicular view of the camera. It should also be mentioned that increasing the on-part resolution is achieved at the cost of reduction in the field of view of the camera. Given a 7 µm/pixel resolution, the field of view of the camera was about 30 mm ×150 mm whereas the field of view in the imaging setup by Kleszczynski et al. , [2], is about 100 mm × 100 mm.Defects such as pores formed in metal powder-bed AM can be as small as 20 µm in size, and pores as small as 100 µm are known to have effect on the part quality and structural properties. With a pixel size of 20 µm, pores of size 100 µm, may appear in only 5 pixels along the diameter.

With a pixel size of 20 µm, pores of size 100 µm, may appear in only 5 pixels along the diameter. Objects of size 5 pixels may easily be produced due to noise, shades, or other non-defective surface features or discolorations. Therefore, for visualization of pores, a smaller pixel size, such as 7 µm, is required which captures a pore of 100 µm in about 15 pixels along the diameter. A 15 pixel object size provides enough visual features for human eyes to identify and distinguish from most other objects. A commercial standard microscope with image pixel size of 7 µm can clearly shows pores of and larger than 100 µm.As seen in Figure 2, the image captured from the layer shows the fusion seams, fused surface and powder bed texture as well as the clear boundary of the fused object. The images provided rather sufficient clarity for segmentation and inspection of the fused region geometry [7].

The images provided rather sufficient clarity for segmentation and inspection of the fused region geometry [7]. However, throughout the fused region, there are several dark regions where the fused surface details are not visible. In fact, they may easily be confused with lack of fusion or defects. A subsequent microscopic examination from the layer, however, showed that these dark areas do not correspond to any defect or pore and are only formed due to shades and insufficient light reflected from those areas.1415Figure 2- Example of image with the setup using 8.8 megapixel camera and adjusted level of intensity for LED spotlight mounted at an angle from the build.To investigate visibility of pores in images, it is required to build layers that contain pores. Therefore, a range of process parameters including non-optimal parameters that are known to produce pores in parts were chosen. Microscopic examination was used to examine and confirm existence of pores in the parts.

Microscopic examination was used to examine and confirm existence of pores in the parts. The target for the visual imaging system in this machine is then to visualize pores as they appear in high-resolution microscopic images.

The resolution of the standard microscope is measured as 7 µm/pixel which turns out to be equal to the camera on-part resolution but with a smaller field of view.From comparison of the camera images by the setup with spot LED as in Figure 1 with microscopic images of the porous layers, it was seen that pores are not visualized in camera image whereas they are clearly visible in microscopic images ( see Figure 3 (a) and (b) for example).(a) (b) (c)Figure 3- Images of the same build layer: (a) captured by a standard microscope with the resolution of 7 µm/pixel, (b) captured by visual camera in the imaging setup using the LED spot light at an angle ( Figure 1 ) with the resolution of 7 µm/pixel, (c) captured by visual camera in the imaging setup using the ring LED perpendicular to the build and around the camera with the resolution of 7 µm/pixel.The reason for invisibility of pores from camera images could be that insufficient illumination is reflected from pores into the camera.

This could be due to large diffusivity and the non-uniform illumination from an angle, as well as the shades of surface elevations that cover the1416pores. More uniform illumination should then be considered that can be provided by implementing two or more light sources from different sides to remove that shadows caused by each individual light source. An alternative more effective illumination would be to use a symmetric illumination such as a ring illumination mounted directly at the top of the build with light arrays perpendicular to the build platform. To provide a uniform, symmetric, and shadow-free illumination such as in microscopes, an adjustable microscope ring light for stereo microscope was used. The ring LED has 2-1/2" (64mm) inside diameter and 4" (100mm) outside diameter. The only place that the ring LED could be mounted, to avoid blockage of the laser beam, was around the camera at the same distance of the camera from the build.

The only place that the ring LED could be mounted, to avoid blockage of the laser beam, was around the camera at the same distance of the camera from the build. The operating distance of similar microscope lights are in the range of several inches, therefore, mounting the LED at this distance of about 1 ft from the part, resulted in dark and rather low-quality images and the problem with capturing dark non- illuminated areas within fused region persisted (see Figure 3 (c)). This problem was most visibly seen within the fused region that has a relatively rough surface and variations of height that can create shadows and prevent the light to be reflected from the surface uniformly. It was investigated and seen that this problem can be removed if the ring illumination is mounted closer to the build at a distance within 5 cm to 10 cm from the build.

It was investigated and seen that this problem can be removed if the ring illumination is mounted closer to the build at a distance within 5 cm to 10 cm from the build. By having light reflected from the build surface uniformly and efficiently, the details of the build surface including surface roughness and height variation as well as pores could be visualized.To mount the ring illumination closer to the build without blocking the laser beam, a ring illumination with a very large inside diameter would have been needed; however, a smaller ring illumination centered closely right on the examined area would be ideal for capturing clear well- lit images. Therefore, it was decided that illumination source would be mounted at a moving fixture so that it stays away as the laser is scanning and moves to the top of the build once the layer is built and is ready to take picture.

Given the current mechanism, it was planned that the ring illumination source be mounted at a fixture connected to the coater and moves away as the coater passes. A fixture with three fixable multi-DoF joints was mounted on the coater (as in Figure 4). The ring LED was mounted on the coater and the fixture was adjusted such that the light would be at the desired distance with the build platform emanating light perpendicular to the platform.

The ring LED was mounted on the coater and the fixture was adjusted such that the light would be at the desired distance with the build platform emanating light perpendicular to the platform. The machine was programmed such that after each layer is scanned or deposited using the laser, the coater would get back and stop at the position where the light is at the specified position and the camera would automatically picture from the layer at that stop.1417Figure 4- Image of the new imaging setup with movable ring LED mounted at a fixture on the coater at a selected height above the build platform.To find an optimal distance for light source from the build platform and adjust the LED at that height, a series of trials and errors by varying the LED distance from the build as well as camera and light settings and parameters were conducted and images of parts were captured.

Camera images were captured in-situ from each layer of several parts built at non-optimal parameters known to form pores. Microscopic images were then captured from the parts’ surface (top layer) outside the machine. It was investigated for the light source mounted at a distance around 60-80 mm from the platform, the camera images captured from the parts show strong correspondence with microscopic images and visualize pores as they appear in microscopic images. This distance was marked and arrangements were planned to mount the light source at that distance from the build.Figure 5 (a) shows an example image captured in situ by the new setup. Figure 5 (b) shows the corresponding microscopic image from the part. It can clearly be seen that the two images match well.

Figure 5 (b) shows the corresponding microscopic image from the part. It can clearly be seen that the two images match well. Additionally the pores are visible in camera images as they appear in microscopic images.1418Figure 5- (a) An example image captured from a porous surface by the setup with movable ring LED (Figure 4), (b) corresponding microscopic image of the part.It was also examined that mounting the camera closer to the build platform would also improve image quality. However, due to the physical limitations and to prevent camera vibration, and since the light source distance from the build had much more dominant effect, it was decided that camera would be remained mounted stably and fixed to the chamber wall.In addition to the distance of the light source from the build, the ring LED is equipped with four zones to independently control light intensity and direction.

The light condition with all four zones lit-up was also compared against only two opposite zones lit-up which the latter resembles the situation with two light source lighting up the part from opposite directions. It was seen the optimal light condition is when all four zones emanate the part uniformly from all directions.In general, this imaging setup is a combination of selected camera settings, use of physical polarizer or filters, illumination intensity and settings, and the distance of the illumination source with the build surface. The most influential parameters are the distance of the illumination source and the camera exposure time; camera gain value and application of filters and polarizer have less important effect. These parameters were adjusted in an iterative manner such that the captured camera images look, to human eye, as much similar to microscopic images as possible. Note that in Figure 5 (a) the color temperature of the camera has been increased (i.e.

Note that in Figure 5 (a) the color temperature of the camera has been increased (i.e. images are red rather than blue as in Figure 2 and Figure 3 (b) and (c)).Although the camera images visualized pores and showed overall good agreements with microscopic images, the camera images still do not show a clear and high-quality picture of the part surface as microscopic images. Further pictures of the surface of the parts were taken after the build and were examined in comparison with their corresponding microscopic images, the reasons for some of the mismatch were later identified and here the main ones can be described as follows: 1) The microscope can be focused very sharply at the top layer within several microns such that extrusions of one layer below are too blurry and are barely seen in the microscopic image.

However, after the camera was focused, it was seen the camera is not focused precisely at the top layer within several microns; in other word, in camera images captured after removal of the powder, extrusions from a layer below were seen as sharp as the features on the top layer (see Figure 6). 2) The high reflectivity of the metallic surface in some regions led to saturations that covered the underneath as well as nearby surface (see Figure 7).1419Lack of a sharp focus on an image in fact acts like a low-pass filter that removes or blurs the sharp and small features associated with low-frequency. Therefore, the color value captured by a pixel is, in fact, affected by the color value (light) from both the location corresponding to that pixel and its neighboring regions. This may cause the images to show some spots darker than they truly are, looking like pores, or slightly distort shapes of some pores.

This may cause the images to show some spots darker than they truly are, looking like pores, or slightly distort shapes of some pores. It may additionally lead to larger saturation regions in the image as pixels receive light from the neighborhood of their corresponding spacial point and pixels that otherwise would not have been saturated, saturate and are considered as missing data (see Figure 7).The two aforementioned problems could be alleviated by further and more detailed modification of camera and ring LED. Overall, current camera images can show individual pores that agree with microscope and the agreement almost holds for any pore larger than 100 microns.(a)(b)Figure 6- Illustration of the sharp focus of camera with respect to microscope. Comparison of the zoomed- in views of image from camera from a part after build (left in each row) and corresponding microscope image (right in each row).

Comparison of the zoomed- in views of image from camera from a part after build (left in each row) and corresponding microscope image (right in each row). Each row shows the same part within the row that is diferent from the other row.1420Figure 7- An example of image in which local image saturation has covered local surface features compared to microscope. Left: camera, right: microscope.Automated Detection of Pores from Camera ImagesAfter the imaging setup enables to capture in-situ images from each layer that visualize detailed surface features and pores, automatic detection of defects can be performed. To detect pores individually in an automatic manner, appropriate image processing algorithms should be developed. The devloped algorithms would have some level of performance in terms of false alarms or false negatives that could be satisfactory or insufficient for a given application or a specific build.

The devloped algorithms would have some level of performance in terms of false alarms or false negatives that could be satisfactory or insufficient for a given application or a specific build. The development and the performance of algoritms would strongly depend on the characteristics and clarity of images, the visibility, contrast, and discriminability of pores compared to the background and non-defective regions, the level of noise, etc.By observing camera and microscopic images of different layers, it is seen that the part surfaces have some discolored regions that appear brown on the surface (see Figure 8). These brown regions are believed to have been formed due to oxidation with the small amounts of Oxygen in the chamber or the gas bubbles in the powder grains [10]. The level of oxygen in the chamber was always mainttained lower than 500 ppm and mainly at 250 ppm.

The level of oxygen in the chamber was always mainttained lower than 500 ppm and mainly at 250 ppm. This oxidation level could significantly be reduced if the build was performed at lower oxygen concentrations such as 100 ppm which is completely feasible. These brown regions, also called darkly etching regions, in images, look like pores in the layers and, unless closely observed, they can sometimes easily be mistaken with pores by visual observation of the images with human eye. Therefore, in processing of the images, they also intefere with segmentation of pores and appear as false alarms.Figure 8- Two example camera images that show a smooth pore-less layer (left) and a porous layer (right). Squares mark examples of darkly-etching regions that are of the same color as pores and sometimes have1421same characteristic of pores after segmentation, and circles mark examples of true pores.

Squares mark examples of darkly-etching regions that are of the same color as pores and sometimes have1421same characteristic of pores after segmentation, and circles mark examples of true pores. Verification was made by microscopic examination.For defect detection, image segmentation algoritms are developed with the objective of detection of pores from images. However, due to the large variety of layer charactersitics, and surface appearances, and different looks of pores, and darkly etching regions, development of an algorihtm to detect individual pores wihout segmentig surface discolorations and darkly etching spots is very challenging and complicated. Therefore, the image segmentation algorithms that are developed to detect pores consequently detect darkly-etched regions as well. After segmentation of pores and darkly-etched regions, the characteristics of the segmented objects can be used to remove many of these false detections.

After segmentation of pores and darkly-etched regions, the characteristics of the segmented objects can be used to remove many of these false detections. Additionally, it will be seen that smoother surfaces (built with higher energy) tend to have more and larger darkly-etched regions compared to rougher and porous surfaces (associated with low energy). This observation would suggest that identification of layers that likely have pores (associated with low energy or Zone III as introduced by Starr et al. [11, 12]) from layers with smoother surfaces that have very low chances of having pores (optimal or high energy, Zones I and II [11]) would then help to more confidently decide if the segmentd objects in the image are pores or darkly etching regions. This identification has been performed and will be briefly presented in the next section.In the following, approprite image procesing algorithms will be developed to detect pores using inetnsity-based algorithms.

Figure 9 shows an example image of a porous layer. As it is seen, there are some darkly-etching spots that have the same color as pores and even sometimes it is difficult for human eye to distinguish between the two. Pores and these dark spots can very clearly be distinguished by directly looking through microscope with both eyes that can detect depth. Microscopic images also in general show relatively good discrimination between pores and these dark spots. However, due to the reasons mentioned in previous section, camera images do not show the level of clarity of microscopic images.Figure 9- Example image of a porous layer1422To start designing segmentation algorithms, we would start with thresholding and would study how pores would appear after implementing different threshold values. Figure 10 show results of thresholding with four different thresholds. A high threshold segments most and larger portions of pores, but it also segments large regions of darkly-etched spots.

Figure 10 show results of thresholding with four different thresholds. A high threshold segments most and larger portions of pores, but it also segments large regions of darkly-etched spots. A low threshold segments fewer darkly-etched spots; however, it doesn’t segment all areas of a pore and does not detect some of the small pores at all. It was examined and decided that a low threshold can then be used to detect locations of pores, and a high threshold gives a better estimate of pores shape and size. In other words, the image with a low threshold not too low to segment noisy spots can locate or mark defects and the image with a higher threshold shows the full shape of defects at the markers. This context can be implemented using an image processing algorithm called reconstruction [13]. Therefore, the image with a low threshold (Figure 10 (b)) was used as a marker to signify locations of pores.

Therefore, the image with a low threshold (Figure 10 (b)) was used as a marker to signify locations of pores. An image with high threshold, Figure 10 (d), was also used as a mask in the (morphological) reconstruction operation.framed imageframed imageframed imageframed imageFigure 10- Result of thresholding of wiener-filtered gray-scale image of Figure 9 with different thresholds. The thresholds are: (a) 0.32, (b) 0.35, (c) 0.4, (d) 0.43.1423framed imageFigure 11- Application of reconstruction to segment pores from the camera image. Figure 10 (b) was used as markerand Figure 10 (d) as mask.Figure 11 shows the result of reconstruction. The result shows that the majority of pores in the layer are successfully segmented with very few false segmentations. Figure 12 shows the result of implementation of the same algorithm to another examples image.

Figure 12 shows the result of implementation of the same algorithm to another examples image. By close visual examination of segmentation results for various images it can be seen that almost all pores larger than 80 microns are segmented in addition to many of the smaller pores. Although applying the same algorithm to other images and layers leads to segmentation of the majority of pores, in some cases, especially towards the higher energy end, several darkly-etching regions are also segmented. It was investigated that it is possible to remove many of these false segmentation by imposing rules on the geometric characteristics of the segmented objects.

It was investigated that it is possible to remove many of these false segmentation by imposing rules on the geometric characteristics of the segmented objects. Identification and removal of these false segmentations and evaluation of the pore segmentation algorithms is among the future work.I18Figure 12- Example of application of the segmentation algorithm to another image of a porous region.Identification of Porous Regions using Intelligent Pattern-Matching Algorithms1424While detection of individual pores using pore segmentation algorithm provides information about layer porosity, it is possible to extract features from the layer that are correlated with porosity and use them in an intelligent identification algorithm to provide a qualitative or even quantitative assessment of porosity. For example, porous layers can be identified versus non- porous layers using such identification.

For example, porous layers can be identified versus non- porous layers using such identification. Additionally, the result of identification of porous regions can serve with detection of individual pores, by suggesting which images represent a porous layer to be detected for pores. To perform identification of porous regions, it is possible to use a discriminator associated with non-porous regions and identify them from porous regions. This identification would also serve as a quantitative measure of porosity in the layers and the part as a whole. This would be possible using application of classification algorithms to identify the texture associated with porous regions from non-porous regions. Classification of porous versus non- porous regions was successfully performed and is briefly described in this section.Figure 13 shows an example image of a smooth, non-porous layer versus a porous layer.

It is seen that the smooth surface is associated with smooth variation of intensity and a porous surface is associated with sharp and abrupt variation. Therefore, a feature associated with the spatial variation of intensity would be able to discriminate these surfaces. By study of the behavior of the images in the spatial and frequency domain after implementation of fast flourier transform, five dominant features were identified that were able to discriminate porous surfaces versus smooth non-porous surfaces.I15(a) (b)Figure 13- Example images of (a) a smooth surface, and (b) a porous surface.After choosing discriminating features, a classifier should be chosen and trained. There are a variety of choices of classifiers such as neural network, support vector machine, k-mean clustering, and statistical classifiers such as Bayesian. For this work, a Bayesian network was developed and trained to classify porous versus non-porous surfaces based on over 100 train samples.

For this work, a Bayesian network was developed and trained to classify porous versus non-porous surfaces based on over 100 train samples. The performance of the algorithm was then evaluated by implementing the classifier to 100 test sample. The Bayesian classifier was able to identify porous versus non-porous layers with over 90% accuracy.ConclusionsThis paper describes the system and the work on development of a new imaging setup with the collaboration of the Edison Welding Institute (EWI). Unlike the previous and existing work on development of imaging setup for powder-bed AM, this system resolves, to a significant extent, the challenges associated with the illumination and captures images that visualize detailed surface1425features such as individual pores as they appear in microscopic images. A movable illumination source is conveniently utilized to illuminate the part surface closely and without blocking the laser power.

A movable illumination source is conveniently utilized to illuminate the part surface closely and without blocking the laser power. Appropriate image processing algorithms were developed to detect pores from the part surface automatically and as precisely as possible. In addition to detection of individual pores, an intelligent classification algorithm, based on Bayesian inference, was developed to identify porous versus non-porous layers. The Bayesian network assigns a probability value that shows the likelihood that the inspected layer is porous. The output value can also be used as a measure of the level of porosity of the layer.

The Bayesian network assigns a probability value that shows the likelihood that the inspected layer is porous. The output value can also be used as a measure of the level of porosity of the layer. The developed identification framework can be used similarly to identify and characterize other part attributes such as quality of fusion, layer roughness, and the level of build energy for the layer.AcknowledgementsThe authors would like to thank the Edison Welding Institute for funding this research andproviding the equipment and support for conducting this work.References1.2.3.4.5.6.7.8.9.10.Craeghs, T., et al. Online quality control of selective laser melting. in 20th Solid Freeform Fabrication (SFF) symposium. 2011. Austin (Texas). Kleszczynski, S., z.J. Joschka, and J.T. Sehrt. Error detection in laser beam melting systems by high resolution imaging. in Proceedings of the Twenty Third Annual International Solid Freeform Fabrication Symposium. 2012.

The analysis primarily focused on evaluating microstructural aspects, defects, and grain size, along with the mechanical properties, specifically the Vickers hardness at various positions within the samples. The findings indicated that the integration of the predefined locally double melting scan in each layer had a significant influence on the microstructure, resulting in variations in grain size across different locations, as well as hardness values with variations of up to 10% across different areas. Moreover, these discoveries underscore the potential of employing the predefined locally double melting strategy in each layer to create fabricated components with distinctive behaviors, like composites, which could find applications in the aerospace industry.Keywords: Laser powder bed fusion, In-situ strengthening, Ti6Al4V.1.

IntroductionTi6Al4V(Ti64) have positioned itself as the most advantageous Titanium alloy in the recent decades due to its demand in the aerospace and the biomedical industries [1-2]. The combination of both α and β phases in the alloy combination makes it more versatile for the varied mechanical properties that can be given rise with a control over the thermal processing strategies [3]. The laser processability of the alloys had paved breakthrough pathways to fabricate complex near net shape parts using Laser Powder Bed Fusion (LPBF). Numerous reviews have been made available discussing the use of the alloy in Additive Manufacturing(AM) [4–6].LPBF has long been used as a successful additive manufacturing process for metals that can bring about micro-sized features at the expense of the limiting laser spot size. A wide variety of alloys are being developed for the process making it a promising AM technique adoption.

A wide variety of alloys are being developed for the process making it a promising AM technique adoption. Therefore, the process have also been developed further to make it time saving for shorted lead time using multiple lasers or area wise processing with the only limitation being the laser spot size [7-8]. On the other hand, the process contributes lower amount of carbon dioxide to the atmosphere than conventional manufacturing process for Titanium alloys [9]. This makes it more advantageous and low carbon footprint process unlike others.1148Post heat treatment is a commonly used after fabrication stage for Ti64 to achieve the desired microstructural phases of the LPBFed part [10]. But it is usually a time-consuming process used right after the fabrication raising the lead time for a finished and processed end part [11].

But it is usually a time-consuming process used right after the fabrication raising the lead time for a finished and processed end part [11]. The conventional heat treatment processes involve raising the temperature of the part and maintaining it at certain time cycles for the desired microstructural phases based on the cooling rate induced. This requires additional cost and time before the final part arrives in hand. The process flow of the LPBF itself paves way for thermal processing by tweaking the laser parameters during fabrication. Since the laser can be controlled over the coordinates of the build plater to bring about the necessary thermal behavior at different regions, the extent of its relationship with strengthening a part layer by layer have been ideated here.

An in-situ microstructural strengthening technique that can bring about desired phases of Titanium drawing parallels with a steel bar strengthening a concrete pillar which could result in a better mechanical property would be a great time and cost saving technique for LPBF. Here, such processing where a strengthening matrix like laser processing on the Ti64 parts has been studied. The results drawn from the study proved that a fabricated part can be strengthened with a refined material in a localized region. This brings about a microstructural based location wise control forming a supporting matrix for the entire part similar to a fiber reinforced composite[12].2. Materials and Methods2.1 Powder and Fabrication ProcessThe powder used for the fabrication of parts in this study, is Grade 5 Ti64 provided by EOS GmbH. Particle size distribution of the powder had been ensured using a sieve size of 80 µm.

Particle size distribution of the powder had been ensured using a sieve size of 80 µm. Based on a prior powder particle distribution analysis using SEM, the particle size ranged between 35 µm and 45 µm. A mix of fresh and reused powder after sieving had been added to dispense for fabrication of the parts using EOS M290. The powder composition as mentioned by the provider is added in table 1.Table 1. Composition of elements in the Ti-6Al-4V powder material purchased from EOS. Element Concentration Balance 5.5 – 6.75TiAlV 3.5 – 4.5 wt.%O < 2000 ppmN < 500 ppmC < 800 ppmH < 150 ppmwt.%2.2 Parts and Fabrication ParametersIn order to study the effect of double exposing on the mechanical property of the fabricated samples, 2 different variations have been studied. One is named DEB which has wider double exposed regions and other one is named as DES which has narrower double exposed regions. The dimension of the coupons used in this study have been shown in the figure.

The dimension of the coupons used in this study have been shown in the figure. The light grey region with dashed arrows is first exposed with laser parameter set A and the dark grey regions with solid1149Fe < 3000 ppmblack arrows are exposed a second time in the coupons with parameter set B. So, the darker regions are exposed twice.Fig. 1: Coupon Dimensions – DEB(Left); DES(Right)The parameter set A is suggested by EOS for the Grade 5 Ti6Al4V powder supplied by them. The set B has been used arbitrarily as 85% reduction in the laser power keeping everything else as is. Set A was processed with a contour using EOS suggested laser parameters for the contour processing and Set B was processed without any contour. So, layer by layer, the sample had been laser processed once with set A and the darker region a second time with set B.

So, layer by layer, the sample had been laser processed once with set A and the darker region a second time with set B. Each of the alternating regions have been named as SE for single exposed and DE for double exposed region with the number indicating its position in the sample.Table 2: Laser Fabrication Parameter sets usedParameter Set A BLaser Power (W) 280 238Scan Speed (mm/s) 1300 1300Hatch Distance (mm) 0.12 0.12Laser Energy (J/mm3) 44.87 38.142.5 Microstructure, Phase and Mechanical Property AnalysisThe region perpendicular to the building direction away from the build plate of the fabricated coupons have been polished until 0.04 µm colloidal silica solution. The samples were then optical micro graphed at various DE and SE regions to study the porosity of the samples at those regions. Image J has been used to measure the 2D surface porosity in different regions of the sample.

Image J has been used to measure the 2D surface porosity in different regions of the sample. The samples have also been studied under Scanning Electron Microscope for phases of the material. The captured images at 800 and 900 magnification level have been used to calculate the lath size distributions using Image J.1150One of the double and single exposed regions from the DEB samples have been cut out from the center to measure the X-ray diffractions due to grain structures formed only due to double exposing and single exposing. The XRD analysis have been performed with Bruker D8 system on both those cut out regions to analyze the effect of the varied laser processing at the lattice level.Vickers Microhardness analysis was used to characterize the mechanical hardness property of the regions in the coupons.

At least 6 indentations were made in each alternative single and double exposed regions of the same sample to capture the hardness variation across the sample perpendicular to the build direction.3. Results and DiscussionThe double exposed regions of the sample can be clearly distinguished from the single exposed regions right out of fabrication by visual inspection. The surface depicted a similarity of that as a carbon reinforced composite where the alternate fiber weave can be visually identified. So, a composite looking like parts has been the resulting parts out of this processing techniques with the mechanical property results supporting it.3.1 Hardness VariationThe micro-hardness indentations made across the specimens reveal that hardness property of a region varied depending on the laser processing conditions in or around the region.

It has been observed in the DEB specimen that the hardness of the coupons had a zigzag trend across the coupons with crests at the double exposed regions and the troughs at the single exposed regions. Whereas the DES specimen showed a collective concentration of the hardness property over a set of alternative single and double exposed regions. The trend observed in the DEB specimen, clearly shows that either microstructural phase strengthening, or a defect density reduction seem to be resulting in an improvement of micro-hardness values in the double exposed regions.

The hardness of the surface perpendicular to the building direction at the region away from the building plate at the sample processed with only set A for reference have also been indicated with a red box marker in fig 2.DE3 SE4 DE4 SE5 DE5Fig 2: Micro-hardness results for DES (left) and DEB (right)1151The micro-hardness indentations made in the coupons reveal that there is a distinguishable variation of hardness values depending on the width of the double exposed regions. The coupons with wider double exposed regions (DEB) seem to have better control over the mechanical hardness of the regions. The narrower double exposed strips give rise to higher heat transfer to the adjacent regions thus contributing to changes in hardness collectively over a set of strips. The width of the double exposed regions was found to be the limiting factor for controlling the hardness.

The width of the double exposed regions was found to be the limiting factor for controlling the hardness. This could be attributed to the defect density or the resulting grain structure as a result varied cooling rates in different regions.3.2. Hardness and PorosityAs mentioned previously, in order to closely study the contributing factor towards increasing hardness the porosity of the regions have been mapped. 2D surface porosity of some of the regions focused to the middle portion of both the DEB and DES samples have been attached in figure. It is observed that as the region's porosity increased, the hardness of those regions also increased. It can be observed that the mechanical hardness property was found to be directly proportional to the defect density in the region. Whereas Javidrad et al. concluded that the defect density affects the hardness with direct proportionality [13].

It has been found that DE regions have finer α lath sizes compared to SE regions. It should also be noted that the double exposed regions had a dark microstructural feature which was not identified in SE region as commonly as in the DE region. The feature has been identified with a red border on the images in Figure 4 (a) and (b). It is believed that such microstructural features lead to finer laths in those regions.1152a)b)0.880.86e)c) ))d))m µ (s s e n k c i h T h t a L0.840.820.80.780.760.740.72DESERegions on DEBFigure 4: a) and b) – SEM image from DE3 region of DEB sample; c) and d) – SEM image from SE4 region of DEB sample; e) Lath Thickness of DE and SE regions3.4. Phase analysisX-ray diffractions have been measured specifically from DE and SE regions cut out from the DEB sample to study the lattice distortions and orientations.

Phase analysisX-ray diffractions have been measured specifically from DE and SE regions cut out from the DEB sample to study the lattice distortions and orientations. DEB sample had been selected due to its strong effect on the resulting hardness over DE regions compared to SE regions. The peaks obtained from the analysis proved that a peak narrowing has occurred after exposing the region twice with set A and set B of laser parameters. This proves that larger crystalline sizes have occurred at the DE region compared to that in the SE region using the Scherrer equation [14]. It can be concluded from the diffraction analysis that, a lattice distortion has occurred in the DE regions because of exposing it twice.700600DE Main Peak) . u500SE Main Peak.

It can be concluded from the diffraction analysis that, a lattice distortion has occurred in the DE regions because of exposing it twice.700600DE Main Peak) . u500SE Main Peak. a (y t i s n e t n I40030020010003939.54040.5 Angle (2Ɵ)4141.542Figure 5 : XRD results from the crystalline phases of a DE and SE region in DEB1153According to the Hall-Petch relationship, the crystalline size is inversely proportional to the hardness of the region [15]. The peak widening in the SE region resulted in a lower hardness value in the region as per figure 2.4. ConclusionThe above explained in-situ laser technique during the LPBF fabrication process have been studied for it effect on the microstructural phases and the resulting mechanical properties. 2D surface porosity, lath thicknesses, lattice distortion and microhardness have been analyzed to find the resulting effect on the fabricated Ti6Al4V part due to the proposed processing technique.

2D surface porosity, lath thicknesses, lattice distortion and microhardness have been analyzed to find the resulting effect on the fabricated Ti6Al4V part due to the proposed processing technique. The highlighting conclusions from the study include:1. The hardness of the regions scanned twice with laser was harder than the regions scanned only once2. Porosity analysis showed that the hardness and porosity were directly related 3. Lath thickness analysis showed that double exposing resulted in finer α laths in the DE regions whereas laths were thicker in the SE regions4. SE regions seem to have higher crystalline sizes compared to that of the DE regions based on the X-ray diffractions analyzed from the lattice structure formed5. AcknowledgmentThis work was supported by a University of Texas System STARs award.6. References[1] L. E. Murr et al., “Microstructure and mechanical behavior of Ti–6Al–4V produced byrapid-layer manufacturing, for biomedical applications,” J.

-Situ Thermal Image Correlation with Mechanical Properties of Nylon-12 in SLS Walker Wroe, Jessica Gladstone, Timothy Phillips, Austin McElroy, Scott Fish, Joseph BeamanThe University of Texas at Austin Mechanical Engineering Department, Cockrell School of Engineering Austin, TX Email: walkerwroe@gmail.comAbstract Selective laser sintering (SLS) of Nylon is a significant portion of the additive manufacturing market for structurally sensitive applications. Current methods of acceptance for such parts are based on the inclusion of ASTM tensile test specimens within the build volume to assess the overall build quality. Ultimate strength and elongation of these specimens oriented both in-plane and normal to the layer build direction are the primary quality metrics.

Ultimate strength and elongation of these specimens oriented both in-plane and normal to the layer build direction are the primary quality metrics. This paper looks at a more complete method of certifying parts for acceptance based on examination of the build conditions in each layer of the part by comparing layer-by-layer thermal conditions during the part build to the resulting ASTM specimen tensile properties. Through such a comparison, a more complete three-dimensional assessment of part quality during the build process can be constructed. The layer-by-layer assessment used here is derived from infrared thermal imaging; mapping temperature profiles of SLS-built tensile bars with data collected before, during, and after each layer-wise laser melting sequence. Mechanical properties and fracture conditions are then quantified and correlated with the conditions where the fractures occur.

Mechanical properties and fracture conditions are then quantified and correlated with the conditions where the fractures occur. Build conditions associated with poor failure conditions may then be used to assess poor SLS bonding throughout the part volume, improving overall part quality assessment and certification. As the method is matured, real time layer-by-layer assessment will be linked to SLS control, to correct for observed defects during the build and improve overall part quality and repeatability. 1. IntroductionAlthough a number of Additive Manufacturing (AM) technologies exist, Selective Laser Sintering (SLS) remains one of the most versatile of these methods for building structurally efficient products. The SLS process creates a three-dimensional part, layer- by-layer, from powder that is melted using a laser. This method is particularly useful for complex geometries without the need (in plastics) for support structures.

This method is particularly useful for complex geometries without the need (in plastics) for support structures. Although SLS has been in use for prototyping and low stress part production since the late 1980’s, further research and development is needed to achieve consistent high mechanical performance in manufacturing applications.For parts manufactured with layer-by-layer accumulation of material (including SLS),the weakest direction is often in the direction perpendicular to the layers, which will be referred to in this paper as the Z-Axis1,2.One important process-based issue in SLS is consistent control of thermal conditionsduring the build. Inconsistent melt and re-solidification conditions can give rise to unpredictable structural results, due to unwanted porosity, and poor bonding between832melt regions both in and between layers. Process monitoring of such issues is essential to understanding their relative importance.

Process monitoring of such issues is essential to understanding their relative importance. [3 (Yuan), 4(Wegner)]The Laser Additive Manufacturing Pilot System (LAMPS) is a valuable SLS researchmachine, developed to easily add or modify measurement and control tools. Visual and infrared (IR) cameras are currently used for in-situ process monitoring as will be described further in this paper. The IR camera is oriented to provide simultaneous thermal measurement of the build surface and powder prior to spreading over the build surface for control of the these temperatures independently from the same image sensor source.With this experimental capability, correlations can be drawn between the thermal history of parts built and their corresponding tensile test results. The focus for this paper on Z-axis tensile bars enables direct comparison of interlayer bonding characteristics under different build thermal conditions with the mechanical properties of ASTM class specimens.

Once this correlation is well characterized, such comparisons can be made (and control compensation implemented) for separate 3D parts.2. LAMPS Machine and Tensile Test SpecimensThe LAMPS (Laser Additive Manufacturing Pilot System) machine was designed and built at the University of Texas at Austin for the study of SLS process control and its extension to existing and new polymer based materials. LAMPS provides direct access to all measurements for both monitoring and control. Controls are implemented in an open architecture using National Instruments LabVIEW based software tools. Users have control over all process parameters, and are able to add and remove data acquisition elements as needed. For this paper, the primary thermal sensor is a long wave infrared (IR) camera (FLIR A325), mounted as shown in Figures 1 and 2.

For this paper, the primary thermal sensor is a long wave infrared (IR) camera (FLIR A325), mounted as shown in Figures 1 and 2. A more complete description of the LAMPS System can be found in Fish et.al.2Figure 1 – LAMPS External View833Figure 2 – LAMPS Build Chamber Cross SectionThe baseline control use of the IR camera is for real time closed loop control of apair of quartz lamps that act as a secondary heating source for the build surface and pre- spread powder drop zones. Background heating of the build is achieved by controlling the atmospheric temperature with an array of wall and build plate mounted heaters. The atmosphere for this experiment is Nitrogen. The IR camera also captures thermal images of the build surface during the laser scan for analysis of thermal conditions of the part layers as they are built, which is the focus of this paper. The images captured for this work at 30 frames per second, over an average scan time of 20 seconds per layer.

The images captured for this work at 30 frames per second, over an average scan time of 20 seconds per layer. The tensile specimens for this paper required 712 layers, requiring over 150 GB of data for the entire build.3. Specimen Build and Corresponding Image AnalysisIn order to perform a direct and consistent comparison between various SLS-built specimens, ASTM Type I tensile specimens with slightly shorter gauge lengths were built as shown in Figure 3. To assess layer-to-layer bonding characteristics, the specimens were oriented perpendicular to the layer plane in 3 rows of 10 specimens each. All specimens were numbered based on their location (row/column) in the build for comparison/correlation with in-situ IR measurements.Figure 3 - Tensile specimen orientation for this study834IR data was captured for each layer of the build for the complete period of laser operation.

Each image was then processed with an image filter [5] to adjust (keystone affect) for the angle offset from the IR camera’s mounting and to align image pixel data with the reference frame of the parts as scanned by the laser. A depiction of the filter’s impact on the IR images is shown below in Figure 4. In this figure, the left image is the raw image taken by the IR camera, while the right image is the adjusted image in MATLAB, which accounts for the laser scan reference frame. Additionally, to reduce the amount of stored data, a cropping tool was used to only include the sintered geometries and their immediate surrounding surface.Figure 4 - Image keystone filteringOnce the filter is applied, appropriate regions of interest (ROI) were designatedcorresponding to the areas sintered for each specimen.

Figure 5 shows the arrangement of these ROI in the three rows of 10 as built in the LAMPS system with row 1 being farthest from the hopper/start of spreading, and closer to the dump side of the build. For each frame of data, an average temperature is determined for each ROI.Figure 5 – Row references in the corrected data imageFigure 6 shows a time-averaged contour map from the corrected imagery for all pixels within a single layer. The elevated humps in this figure correspond to the ROI’s for each specimen, and one can see that the background temperature is higher above the 3rd row.

The elevated humps in this figure correspond to the ROI’s for each specimen, and one can see that the background temperature is higher above the 3rd row. This perceived higher background temperature is likely due to radiation bleed835around a shield between one of the quartz lamps and the IR camera and may be biasing the ROI derived temperatures for the 3rd row of specimens.Figure 6 - Build surface temperatures averaged over one layer scan periodAfter all images were captured for a single layer, a temperature matrix was built for the array of specimens corresponding to several key conditions within a layer. The first is an average temperature for each ROI over the complete scan time (as shown in Figure ** above).

The first is an average temperature for each ROI over the complete scan time (as shown in Figure ** above). We also extracted a representative temperature for each ROI just before it was scanned and just after it’s scan exposure was completed as described in Section 4.During the build process, the laser scanned in alternating directions by layernumber as is common for commercial scanning practice, as shown in Figure 7. For odd layers, starting at Layer 1, the laser scans or “marks” in the positive x-direction, and then jumps back and repeats this movement at the next higher value of y based on the scan separation distance. For even layers, the laser scans in the positive y-direction and then repeats with the scan offset increasing in x.836Figure 7 - Laser scanning pattern by layer number4.

For even layers, the laser scans in the positive y-direction and then repeats with the scan offset increasing in x.836Figure 7 - Laser scanning pattern by layer number4. Description of Thermal Imaging AnalysisIn addition to the average temperature of each specimen’s layer over the complete scan period, an estimate of the temperature just prior to scan and just after the scan is also desired for correlation with observed mechanical properties. The image frames within a layer for these conditions were found as follows. First, the frame-by-frame temperature of each specimen over the layer scan period is gathered. Examples of this temporal variation for two specimens are plotted in Figure 8 for an odd layer. ROI 1 corresponds to the first specimen in the 1st row and ROI 11 corresponds to the first specimen in the 2nd row (See previous ROI Figure 5 for reference).

ROI 1 corresponds to the first specimen in the 1st row and ROI 11 corresponds to the first specimen in the 2nd row (See previous ROI Figure 5 for reference). When an image frame captures the laser scanning the ROI of interest, the associated average temperature for that ROI jumps out of range due to the very high reflection in the IR band for the CO2 laser being used. One can nevertheless follow the effective average temperature by ignoring these frames as illustrated in Figure 8. The frame where and ROI first jumps is a good indication of when scanning starts for that specimen and the previous frame is used as its “pre-scan” temperature.

The frame where and ROI first jumps is a good indication of when scanning starts for that specimen and the previous frame is used as its “pre-scan” temperature. For odd layers, each row is effectively finished scanning when the next row scanning starts, so the frame taken when ROI 11 temperature jumps out of range (in row 2) is selected for evaluation of “post scan” temperatures for all specimens in row 1.Figure 8 - Sample image frame based temperature history for two specimens (ROI) for a single layer scan.837Pre and post scan temperatures were thus gathered for all specimens over alllayers of the build.5. Mechanical and Thermal ResultsAll thirty specimens were tensile tested using standard ASTM 638 methodology using an Instron 3345 tensile machine and the resulting stress vs. strain curves for each are shown in Figure 9. All specimens fractured within the gauge length, except for the weakest specimen, which fractured in two places simultaneously.

strain curves for each are shown in Figure 9. All specimens fractured within the gauge length, except for the weakest specimen, which fractured in two places simultaneously. The near vertical lines at maximum extension aid in visualizing the strain at break for each specimen.Figure 9 - Measured stress vs. strain for all 30 tensile specimensThe weakest specimens corresponded to the lower average temperature regions in the build, which were located on the ends of each row as shown in Figure 5, from Section 3. A more thorough study of spatial variation in specimens will be covered in a future paper. The first comparison uses the average temperature over the full scan time of all layers and the corresponding fracture strength as shown in Figure 10. Other than the two specimens at 185.1 and 185.5 C, there appears to be a break in the strength behavior of the group at around 186.5 C.

Other than the two specimens at 185.1 and 185.5 C, there appears to be a break in the strength behavior of the group at around 186.5 C. Linear fits to the data (excluding the two points mentioned above) above and below this break are shown suggest that the fracture strength drops off by 75% or more if the average temperature does not reach the break point. Further discussion of this can be found in Section 6. A significant increase in strength, as the average temperature is increased, is seen in the higher strength portions of Figure 10.838Figure 10 - Specimen “scan time” average temperature vs. fracture stressFigures 11 and 12 show the similar correlation for pre-sintering and post sintering temperatures averaged over all layers, for each specimen versus its fracture strength. The correlation with pre sintering is obviously poor, while the post sintering temperature shows a two-zone characteristic similar to that with the overall average layer temperature described previously.

The post sintering temperature, which should be a more direct measurement of layer melt and bonding quality, does not show as strong a correlation with fracture strength as the scan averaged data of Figure 10, and also shows significant variation in strength around the transition temperature between low and high strength specimens.Figure 11 - Averaged pre-sintering temperature vs. fracture stress839Figure 12 - Averaged post-sintering temperature vs. fracture stressCorrelations between the strain at break and the pre- and post- sintering layeraveraged temperatures are shown Figures 13 and 14, respectively, and follow the nature of the strength data.Figure 13 - Average pre-sintering temperature vs. strain at break840Figure 14 - Average post-sintering temperature vs.

strain at break840Figure 14 - Average post-sintering temperature vs. strain at breakAlthough the trend is not as obvious as that of the tensile stress plots, it is stillclear that the post-sintering temperature of a specimen plays a bigger role in the maximum tensile strain. For low post-sintering temperatures, the tensile strain is only 4- 7%, while at higher post-sintering temperatures; the strain can nearly be doubled, at about 10-12%.It is believed that the reason the above correlations are not very strong is that byaveraging over an entire ROI (35 pixels), the initially local nature of delamination between layers can be washed out. In order to identify these local defects a new analysis strategy was implemented where the minimum average temperature of every 3x3 pixel area in each ROI was taken from the post-sintering image and chosen as the temperature to represent that layer.

The minimum value of all the layers was then taken to represent the temperature value for that specimen and is called the local-minimum post-sintering temperature. Figure 15 below gives a visual description of the 3x3 regions in each ROI, and Figure 16 shows the plot of the local-minimum post-sintering temperature versus ultimate tensile strength. This new comparison gives a much stronger correlation of .746, with the correlation between the local-minimum post-sintering temperature and tensile strain at break being .724.841Figure 15 – Finding local-minimum post-sintering temperatureFigure 16 – Local-minimum post-sintering vs. ultimate tensile strength6. Examination of Specimen Fracture CharacteristicsAdditionally, due to the poor layer-to-layer bonding experienced in z-directionbuilds, the specimens all notably broke with brittle characteristics and well aligned with the layer orientations. Figure 17 shows sample images of a strong and weak break, respectively.

Figure 17 shows sample images of a strong and weak break, respectively. It is difficult at this scale to see differences in the break characteristics but842we expect a closer examination would reveal characteristics of unmelted power in the lower strength samples.Figure 17 - Sample fracture images for low and high strength specimens7. Localized Layer Temperature vs. Fracture CorrelationAll sample fracture locations were measured and linked to a build layer number. This was accomplished by measuring the distance from the bottom of the specimen to the fracture with an uncertainty of +/- 2.5 mm estimated from material ejection during fracture and caliper accuracy. This corresponds to a layer uncertainty band of +/- 30 layers in the build. Figures 18 and 19 show examples of both positive and negative correlation between thermal and mechanical layer failure criteria performed for each specimen using the method described below.

Figures 18 and 19 show examples of both positive and negative correlation between thermal and mechanical layer failure criteria performed for each specimen using the method described below. The process begins by taking each layer’s average temperature and non-dimensionalizing it as a percent difference from the overall layer mean value. Note that the raw layer temperature data are clumped either above or below the mean temperature with few being close to the mean value. This difference in temperature is a result of increased energy deposition by the laser at the beginning and end of each “mark” through a specimen.

This difference in temperature is a result of increased energy deposition by the laser at the beginning and end of each “mark” through a specimen. Because odd layers have this increase over the long sides of the rectangular area of each specimen, they have a slightly higher temperature over the ROI than for even layers which have the narrow edges getting more laser energy.A low pass filter with a pass-band frequency of 0.1, corresponding to a blending of 10 layers, mitigates this oscillation and captures general trends in the temperatures of the specimen as a function of layer. In the figures, individual layer data is shown with circles and the filtered data shown with a solid curve. The different energy density levels in even and odd layers need further investigation, but an initial compensation has been examined with results shown in Figures 20 and 21 for the same data as shown in Figures 18 and 19.

In this new view, odd and even layer deviations are shown from their respective mean values. This reduces the layer-to-layer oscillation and allows for a higher frequency low pass filter of 0.25, corresponding to blending of 4 layers, and a more direct examination of unusually cool layer temperatures compared with their averages. From the filtered thermal layer data, we extract the layer number where the lowest ten-layer blended temperature occurs, which is shown as the red circle icon on the curve. This is based on an assumption that failure will occur close to the layers with the lowest collective melting and bonding. This lowest temperature layer is then compared with the band of potential failure layers from the mechanical measurement, shown in yellow band in the figure.

This lowest temperature layer is then compared with the band of potential failure layers from the mechanical measurement, shown in yellow band in the figure. If the thermal estimate of weakest layer falls within the band of mechanical failure measurement, it is considered potentially correlated.843Figure 18 - Potentially correlated thermal and mechanical layer failure conditionsFigure 19 – Non-correlated thermal and mechanical layer failure conditions844Figure 20 - Potentially correlated thermal and mechanical layer failure conditionsFigure 21 - Non-correlated thermal and mechanical layer failure conditionsUsing these criteria, 12 out of 26 specimens were correlated in thermal andmechanical layer failure, a success rate of 46%, while randomly choosing points for suspected break location results in a success rate of 20%.

Although this percentage is not ideal, it suggests that further examination of the criteria and improved measurement could lead to useable correlation of thermal conditions to break likelihood.The local-minimum post-sintering temperature for each layer was also used to predict the break location but the success rate fell to 35 % against an expected random success rate of 16 %. This decreases in the random predicted success rate is due to a845tightening of the success band, the yellow region in the above figures, from +/- 30 layers to +/- 20 layers.8. ConclusionIn-situ measurement of the thermal properties of the laser sintered surfacesthroughout the build process was analyzed and compared with the mechanical failure stress and elongation of resulting tensile specimens. The results show correlation between the average temperature of the sintered layers and the strength and elongation with a threshold for this average temperature, below which poor performance is noted.

Pre-sintering layer temperatures were not found to correlate well with fracture strength or strain at break. Correlation of fracture strength and elongation was not found with the pre-sintering layer temperature, and the correlation with post sintering temperatures was surprisingly low, given the overall average temperature correlation. When more local temperatures where looked at the correlations did improve significantly to about .74. Initial study of targeting weaker layer regions based on lower thermal measurement looks promising, improved quality of laser scanning control is needed to provide predictive consistency.Future work to improve this prediction will involve closer control of laser powerat the beginning and end of “marking” vectors. This can be accomplished by working with a laser with a more uniform laser energy density.

This can be accomplished by working with a laser with a more uniform laser energy density. Additionally, increasing the time resolution during the scan to better resolve the average temperatures of interest can make improvements to the results. Closer examination of the fracture samples will also be used to better document post fracture inter-layer conditions for comparison with thermal data from the build in the local region of failure.AcknowledgmentThis work is sponsored by the Air Force Research Lab and General Dynamics Information Technology (GDIT) under GDIT Prime Contract FA8650-11-D-5702 / 0009.Citations 1. Yuan, M., and D. Bourell. "Efforts to Reduce Part Bed Thermal Gradients during Laser Sintering Processing." Proc. of 23rd Annual International Solid Freeform Fabrication Symposium, Austin, Tx. N.p.: n.p., n.d. N. pag. Print.2. Fish, S. S. Kubiak, A. Bryant, J. Beaman, W. Wroe, and J.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIn-Situ Thermographic Inspection for Laser Powder Bed FusionT. Liu1, C.S. Lough1, H. Sehhat1, J. Huang2, E.C. Kinzel3, M.C. Leu11Mechanical and Aerospace Engineering Department, Missouri University of Science and Technology, Rolla, MO 65409, USA2Electrical and Computer Engineering Department, Missouri University of Science and Technology, Rolla, MO 65409, USA3Aerospace and Mechanical Engineering Department, University of Notre Dame, Notre Dame, IN 46556, USAAbstractLaser powder bed fusion is strongly influenced by the quality of the powder layer including the powder properties as well as the layer thickness. In particular, thermal stresses can produce sufficient part deformation to the point that a part interferes with the wiper.

In particular, thermal stresses can produce sufficient part deformation to the point that a part interferes with the wiper. This paper investigates the use of long-wave infrared thermography to monitor the surface temperature of the build. When cold powder is spread by the wiper, heat diffuses from the underlying part through the powder. The surface temperature history is a strong function of the thermal transport properties of the powder as well as the thickness. This correlation is explored and measured experimentally. It is then used to estimate the powder layer thickness above overhanging parts. This approach is shown to capture the part distortion and predicts wiper state prior to catastrophic interaction with the part.Keywords: Laser powder bed fusion, long wavelength infrared, powder thickness1.

IntroductionLaser powder bed fusion (LPBF), as one of the most common additive manufacturing technologies, is a process that melts a metal powder by a high-power laser beam layer by layer to generate 3D metal parts [1-3]. The powder layer properties affect the product performance significantly. For example, the microhardness decreases, and the grain size increases for 1Cr18Ni9Ti stainless steel powder when the layer thickness increases in the range of 60 – 100 µm [4]. However, for the CL50WS powder, powder layer thickness in the range of 30 - 40 µm has no significant influence on microhardness [5-6]. Other mechanical properties of the fabricated parts may also be sensitive to powder properties. For the parts produced from Ti-6Al-4V powder, rapid increase in porosity and surface roughness may result from the use of thick powder layer above 40 µm [7].In-situ measurement of powder layer thickness has attracted many researchers’ attention in recent years. Richard et al.

Richard et al. developed a laser displacement in-situ non-destructive method to measure the top powder layer thickness [8]. There are other optical in-situ methods to measure the top powder surface profile. For example, the high-speed spectral domain optical coherence tomography (SD-OCT) was used to measure the layer roughness [9]. The digital fringe projection technique was implemented to measure the surface topography of powder layer [10]. Although these optical approaches can measure the powder layer thickness, none of them can provide the powder thermal conductivity information based on their in-situ measurement data.Thermal cameras and imaging approaches have been used to monitor thermal features during LPBF manufacturing [11]. Short-wavelength infrared (SWIR) cameras, which are more sensitive at high temperature, have been used to monitor the melt-pool dynamics and stability [12], cooling rate [13], and combined with visible cameras to monitor defects [14].

In contrast, relatively inexpensive, uncooled long-wavelength infrared (LWIR) cameras, which are more sensitive at lower temperature, have been used to observe the plume formed by material evaporation and heating of the surrounding gas [15], obtain the surface temperature and cooling time [16], and predict the temperature distribution [17]. Furthermore, image data has been combined with machine learning algorithms to provide an efficient means to assess part quality, validate specific internal geometries, and detect build defects [18,19]. However, no previous researchers have used LWIR cameras to measure in-situ the powder layer thickness and thermal conductivity. The study described in the present paper fills this significant gap.In this study, a LWIR camera is used to monitor the temperature of the powder bed during the LPBF process.

The study described in the present paper fills this significant gap.In this study, a LWIR camera is used to monitor the temperature of the powder bed during the LPBF process. By analyzing the thermography data, the temperature profile is calculated from the thermal history, and the temperature rise time is mapped to the powder layer thickness based on 1D thermal modeling. The obtained powder297layer thickness has been verified experimentally. Finally, an application of the in-situ thermographic monitoring for recoater jam prediction is demonstrated.2. Experimental setupRenishaw AM250, which is an additive manufacturing machine that fabricates metal parts by using the LPBF process, was used to manufacture 304L stainless steel samples. An FLIR A655sc LWIR camera was installed in a Renishaw AM250 machine at the Missouri University of Science and Technology. Figure 1a shows the experimental setup.

An FLIR A655sc LWIR camera was installed in a Renishaw AM250 machine at the Missouri University of Science and Technology. Figure 1a shows the experimental setup. The ZnSe window locates in front of the LWIR camera was used to seal the build chamber. Figure 1b shows an example of the camera field of view.(a)Laser Optics Housing(b)LWIR CameraBuild ChamberZnSe WindowBuild Plate25 mmFig. 1. (a) LWIR image of blackbody during cooling and (b) reference visible image of experimental setup.Since the 304L stainless steel powder bed and LPBF printed solid part have different emissivity, the radiance recorded by the infrared camera needs to be calibrated for both the powder bed and the solid part. Our calibration experiment is shown in Fig. 2. A LPBF 304L stainless steel bin with a cavity in it is manufactured to the equivalent of a blackbody. 304L powder was placed partially on the bin’s top surface, and the bin surface remained observable from the infrared camera.

304L powder was placed partially on the bin’s top surface, and the bin surface remained observable from the infrared camera. This blackbody was heated by the AM250 laser with 200W for seconds at a time as shown in Fig. 2a schematic. This design configuration permitted the simultaneous temperature calibration for the blackbody, 304L stainless steel powder, and 304L solid bin. The LWIR image shown in Fig. 2a denotes the temperature distribution during cooling, and a visible image of the experimental setup is shown in Fig. 2b. The Tapparent is the temperature measured by the thermal camera single pixel. The actual temperature Tactual of the blackbody was recorded by thermocouples located at the top and bottom of the calibration cube. The average of the two thermocouple values is plotted as the actual temperature in the heating and cooling curve shown in Fig. 3(a). The corresponding Tapparent temperatures for the various material states were measured with the LWIR camera.

3(a). The corresponding Tapparent temperatures for the various material states were measured with the LWIR camera. These results are included in Fig. 3(b).In the cooling process, the various material states have the same temperature. However, the different emissivity of these various states make the apparent temperatures recorded by the camera different as shown in Fig. 3(b). The cooling portions of the curves in Fig. 3(b) were used to develop the calibrations from the apparent temperature to the actual temperature. These results are included as Fig. 4 and show that the actual temperature scales linearly with the apparent temperature for all the blackbody, the 304L powder, and the 304L solid.

3. Blackbody heating and cooling curves with (a) actual temperature from thermocouple measurements and (b) apparent temperature from various locations in LWIR data.800700600]C ° [500la u400t c a T300200Blackbody R2 = 0.99 304L Powder R2 = 0.99 304L Solid R2 = 0.99100002550100 75 Tapparent [°C]125150175Fig. 4. LWIR temperature calibrations for blackbody material, 304L stainless steel powder, and 304L stainless steel solid.299100(b3. LWIR camera data processing3.1 Thermal historyThe specimens of 304L stainless steel were manufactured by the Renishaw AM250 and recorded by layer- to-layer LWIR imaging. The laser in Renishaw AM250 scans the powder parts with a point-to-point exposure strategy. This scanning process steps discretely with the parameters of point distance, exposure time, and hatch spacing.

This scanning process steps discretely with the parameters of point distance, exposure time, and hatch spacing. The laser scanning has a fixed orientation perpendicular to wiper moving direction and there was no rotation of the scanning direction from layer to layer. The wiper is located in front of the build plane during the laser scanning fabrication. After the laser finishes the scanning of one layer, the wiper returns to the powder delivery hopper at the back side of the build plane. Then the wiper obtains the fresh powder from the powder delivery hopper, and it goes forward to spread a fresh layer of powder. This process is shown in Fig. 5. The experiment parameter sets are laser power 200W, exposure time 50 μs, point distance 60 μm, hatch spacing 85 μm, and powder thickness 50 μm. The build plate is heated to 80°C during the LPBF process.

The build plate is heated to 80°C during the LPBF process. The entire build chamber is filled with Argon gas to avoid oxidation.Powder delivery hopperInfrared cameravδWiperPowderPowderPrinted solidPowderBackBuild planeFrontFig. 5. Schematic of powder spreading process The thermal histories of laser printing with the calibration equations in Eq. (1) are shown in Fig. 6a. When the build plate’s top surface is covered by powder, the thermal history needs to be calibrated using the powder emissivity. Similarly, when the powder is melt and then solidified, the solid emissivity should be used. Fig. 6b shows this calibration results. In Fig. 6a, the left-valley minimum temperature indicates the wiper temperature when the wiper is spreading the powder. When the wiper obtains the fresh powder from the powder delivery hopper, it moves to the front to spread a fresh layer of powder with 50 µm height. The newly spread powder is heated by the printed layer underneath.

The newly spread powder is heated by the printed layer underneath. The temperature rise curve after the left valley represents this temperature heat-up process. After this spread, the wiper stays at the front of build plate during the laser printing. The temperature at the peak represents the laser heating the powder bed and melting the powder. After the laser printing, the wiper goes back to the powder delivery hopper. The right-valley minimum temperature in the thermal history indicates the wiper temperature when the powder is covered by the wiper during its traveling back. There is a slight temperature rise when wiper return.

There is a slight temperature rise when wiper return. This is because when wiper return, it respread tiny powder on the top of printed solid area.(a)]C ° [ e r u t a r e p m e T600400200Wiper spreading powderSolid calibration Powder calibration Blackbody calibration TapparentWiper return(b)]C ° [ e r u t a r e p m e T600400200SolidPowderWiperSolidSolidWiper00048 Time [s]1216036 Time [s]9Fig. 6. (a) Temperature history obtained by using different calibration equations, (b) calibrated temperature history.30012The powder temperature heat-up process is valuable because it relates to the powder properties: thetemperature of a thinner powder layer rises faster than the temperature of a thicker powder layer.3.2 1D heat transfer modelIn order to understand the effect of powder properties, a one-dimensional heat transfer model is established. A schematic of the 1D heat transfer model is shown in Fig. 7a.

A schematic of the 1D heat transfer model is shown in Fig. 7a. There are two finite slabs with thickness δ and s representing the powder and printed solid parts. The powder slab is assumed in perfect contact with the printed solid slab. The x-coordinates xp and xs for the powder and solid slabs are conveniently chosen as shown in the figure.

The powder slab is assumed in perfect contact with the printed solid slab. The x-coordinates xp and xs for the powder and solid slabs are conveniently chosen as shown in the figure. The equation describing unsteady 1D heat conduction and the initial and boundary conditions for the powder slab in xp- coordinate are given by∂ T 1 ∂ t=kp C ρ pp2 ∂ T 1 2 ∂ x p( , T δ t 1)=( ) T t ikp∂ T 1 ∂ xp x=0= h T  1(0,t)−T ∞ ( T x 1, 0p)T=pwhere T1 is the temperature distribution of powder, t is time, xp is the distance normal to the build plane, kp is the thermal conductivity of powder, Cp is the specific heat capacity of powder, ρp is the density of powder, δ is the thickness of powder layer, Ti is the interface temperature, T∞ is the ambient temperature, h is the convective heat transfer coefficient of gas, Tp is the initial temperature of powder.

The radiation loss occurred at the powder top surface is assumed negligible.The finite difference method is used to solve the above equations for the temperature distribution for a given powder thickness. The obtained temperature distributions for several powder thicknesses are shown in Fig. 7b and 7c, where the material properties used are given in Table 1. The T∞ is 30°C as the temperature of the Argon gas. Ts is 80°C as the build plate heating temperature. Tp is 30°C as same as T∞, h is 15 W/m2K, s is 3mm.Table 1 Material properties of 304L solid and powderProperties Density ρ [kg/m3] Specific heat capacity c [J/kg•K] Therma conductivity k [W/m•K]304L Solid 304L Powder8030 490 16.24818 490 0.13Figure 7b shows the temperature history of the powder top surface for different powder thicknesses. The temperature of a thicker powder layer rises more slowly than the temperature of a thinner powder layer. Fig. 7c shows the temperature distribution along the depth direction.

7. (a) Schematic image of 1D heat transfer model, (b) top surface thermal history (c) and temperature distribution at various time instants.4. Correlation of thermal features with powder properties4.1 Rise timeThe heat-up curves in Fig. 7b are examined carefully in order to express them in mathematical equations. It is observed that a liner combination of a constant and two exponential functions with different coefficients can be used to fit the temperature rise very well. The expression of this fitting function is−α t 1−α t 2 − A A e 1 2 − A A 1 2A e 3 A t 3 This is a piecewise function. A1 is the powder’s steady-state temperature. A1-A2-A3 is the powder’s initial temperature. α1 can be very large to represent the initial rapid heat-up process and α2 is smaller to represent the speed of approaching the steady-state temperature. ti is the time that the powder starts to heat up. By using this function, the temperature rise curves in Fig.

ti is the time that the powder starts to heat up. By using this function, the temperature rise curves in Fig. 7b can be fitted very well as shown in Fig. 8a. The purpose of constructing this fitting function is to obtain the temperature rise time. The temperature rise time is defined here as the time period from the initial time to the time when temperature reaches 90% of the steady-state temperature. The rise time can be obtained by solving the equation:−t>tiT=  −<t iA e 2−α t 1+A e 3−α t 2=0.1(A 2+A 3)(a)(b)600]C ° [ e r u t a r e p m e T807060504030Various 1D model δ [μm]: 50 100 150 200Fitting Fitting Fitting Fitting]m μ[s s e n k c h tir e d w o P4002001D model Fitting by00.00.20.40.60.81.0024Time [s]Rise time [s]Fig. 8.

8. (a) Comparison between 1D model thermal history with fitting result and (b) powder thickness with respect to rise time calculated from 1D model.302(6)(7)The rise time of different powder thickness calculated by the 1D model is shown in Fig. 8b with respect to the powder thickness. It is clear that the relationship between powder thickness and temperature rise time is monotonic, and it can be represented by the following exponential function:δ=0.5045231.2 risetBy using this exponential function, the powder layer thickness can be calculated from its temperature rise time.4.2 Powder propertiesEquation 8 shows the relationship between powder thickness and temperature rise time calculated by using 1D heat transfer model. This relationship needs to be validated experimentally. The specimens with different heights were fabricated by the AM250 machine for the validation experiment. As shown in Fig.

This relationship needs to be validated experimentally. The specimens with different heights were fabricated by the AM250 machine for the validation experiment. As shown in Fig. 10, there are 4 identical specimens, each having 8 strips with different heights. Fig. 9a shows the profile of 8 different heights and Table 2 lists the strip heights and the corresponding powder thickness for every strip. After this experiment, the build plate was taken out of the build chamber and its precise dimensions was measured by laser scanning. The powder layer thickness above every strip can be calculated as the total powder thickness subtracts the specimen’s height. The 2D map of powder layer thickness is shown in Fig.

The powder layer thickness above every strip can be calculated as the total powder thickness subtracts the specimen’s height. The 2D map of powder layer thickness is shown in Fig. 9c.Table 2 The height of each of the eight different strips and corresponding powder thicknessn h [mm] 1 5.00 2 4.75 3 4.65 4 4.90 5 4.70 6 4.85 7 4.80 8 4.95δ [µm] 50 300 400 150 350 200 250 100To measure the temperature rise, when the manufacture of these 4 specimens was done and before moved it out, the build plate was kept in the Renishaw AM250 build chamber’s original position. Then the 4 specimens’ top surfaces were cleaned. After cleaning specimens’ surface powder and keeping the build plate temperature at 80°C, a new fresh layer was spread by the wiper to cover all the 4 specimens’ top surfaces. Since the fresh powder has a lower temperature Tp and the specimens have different heights, the specimen strips with higher heights have thinner fresh powder to heat up.

Since the fresh powder has a lower temperature Tp and the specimens have different heights, the specimen strips with higher heights have thinner fresh powder to heat up. The fresh powder was heated up at different speeds at various locations. The infrared camera recorded the process of temperature rise for all the height strips. Then Eq. 6 was used to fit all the temperature history and Eq. 7 was used to calculate the rise time. The final 2D map of the rise time is shown in Fig. 10a.Figure 11a shows the temperature rise curves at four different positions on a specimen’s top surface. The fitting curves obtained by Eq. 6 match very well with the temperatures measured by the LWIR camera. The data of mean powder thicknesses vs. temperature rise time for each of the 8 strips in the four specimens from Fig. 9c and 10a are plotted in Fig. 11b. The error bars indicate the standard deviation.

temperature rise time for each of the 8 strips in the four specimens from Fig. 9c and 10a are plotted in Fig. 11b. The error bars indicate the standard deviation. The power function obtained to fit these mean values by regression isδ=t 218.60.58 rise.Note that the values of coefficients in this function from experiment are close to the values of coefficients obtained previously 1D model for Eq. 8. The difference is caused by the rise time calculation. The 2D map of powder thickness can be calculated from Eq. 9 and is shown in Fig. 10b.In order to compare the powder thickness predicted by Eq. 9 with measured powder thickness, data of Fig.9c and Fig. 10b are plotted together as show in Fig. 11. The R2 for this prediction results is 0.7641.303(8)(9)(a)h12δ345678Powder surfaceSpecimen(b)Specimen aSpecimen bSpecimen cSpecimen d10 mmδ [mm](c)0.50Fig. 9.

11. (a) Comparison between measured temperatures with fitting results and (b) powder thickness with respect to temperature rise time from experimental data.3045]m μ500[s s e n k c hit400300R2 = 0.7641r e d w o p d e t c d e r Pi2001000 0100200300400500Powder thickness [µm]Fig. 12. Comparison between measured powder thickness with respect to powder thickness predicted by rise time5. Application of the in-situ thermographic monitoring for recoater jam predictionThe thermographic inspection of powder thickness is applied for recoater jam prediction and prevention in the LPBF process. Because of the residual stresses occurred in the parts during LPBF manufacturing, overhanging parts always curl up. The deformation grows faster on the overhanging part’s contour than its interior. The part curl- up after deformation over a threshold can hit the wiper and cause serious consequences.

The deformation grows faster on the overhanging part’s contour than its interior. The part curl- up after deformation over a threshold can hit the wiper and cause serious consequences. Hence, prediction and prevention of the recoater jam in real-time LPBF manufacture is necessary. A recoater jam prediction system has been established based on calculating the powder layer thickness at different locations of the powder bed. When the curl-up deformation grows, the powder layer at the curl-up location becomes thinner and thinner and eventually will damage the wiper and cause the recoater jam. The powder layer thickness can be used to represent the amount of curl- up. To demonstrate this application, overhanging semi-arches are manufactured, and the corresponding powder thicknesses are calculated. The nominal layer thickness is 50 µm. Figure 13a shows the calculated powder thickness at the right edge of each of two semi-arches shown from the first layer to the occurrence of recoater jam.

The nominal layer thickness is 50 µm. Figure 13a shows the calculated powder thickness at the right edge of each of two semi-arches shown from the first layer to the occurrence of recoater jam. The curve has four different regions. In the initial region shown by the light blue area, the powder thickness increases. In the green region from layer 50 to layer 170, the powder thickness keeps essentially constant. In the yellow region, the powder thickness drops (decreases quickly) from layer 170 to layer 215. In the pink (last) region, the recoater jam occurred and powder thickness fluctuates wildly. Fig. 13b shows the final states of five semi-arches and the ablated wiper in the experiment.

In the pink (last) region, the recoater jam occurred and powder thickness fluctuates wildly. Fig. 13b shows the final states of five semi-arches and the ablated wiper in the experiment. The silicone part of the wiper is ablated by the semi-arch deformations, and the metal parts of wiper behind silicone are scratched too.(a)]m μ150(b)[s s e n k c hitr e d w o p d e a u c a Ctll1209060301Arch 1 Arch 22Metal SiliconeMetalScratch1mm04080120 Layer160200240Ablated wiperFig. 13. (a) Real-time calculated powder thickness and (b) final semi-arch samples and the ablated wiper.6. Conclusion3051mmThis paper presents using a LWIR camera to perform in-situ powder layer thickness during metal additive manufacturing processes. A 1D powder layer heat transfer model was developed and solved using the finite difference method, from which the calculation of rise time profile from temperature distribution in the powder was made.

The method of mapping temperature rise time to powder layer thickness was illustrated. The 1D model of temperature rise time vs. powder thickness agrees well with validation experimental result. The application of this thermographic inspection method for overhang curl-up detection and recoater jam prevention was demonstrated.7. AcknowledgementsThe authors would like to thank Honeywell for its leading role on this research project and UCLA for the cooperative efforts. The ﬁnancial support for this research is provided by the Clean Energy Smart Manufacturing Innovation Institute (CESMII).8. References1. Lough, Cody S., Xin Wang, Christopher C. Smith, Robert G. Landers, Douglas A. Bristow, James A. Drallmeier, Ben Brown, and Edward C. Kinzel. "Correlation of SWIR imaging with LPBF 304L stainless steel part properties." Additive Manufacturing 35 (2020): 101359.2. Hung, Chia-Hung, Wei-Ting Chen, M. Hossein Sehhat, and Ming C. Leu.

Selective Laser Melting (SLM), an additive manufacturing process, of as-prepared TiC/316L composite powder was performed. It was found that the relative density, microstructure and mechanical properties of the SLM-processed nanocomposites were sensitive to the applied energy density. In general, the improvements in the mechanical properties are attributed to the combined effects of grain refinement and grain boundary strengthening. We believe that the alternative fabrication route presented in this study should significantly increase the use of nanocomposites.Keywords: mechanical alloying; selective laser melting; porosity; hardness; wearIntroduction316L stainless steel alloy is of great interest in the industry for their superior ductility and excellent corrosion resistance [1]. However, the applications of these alloys have been limited due to their poor strength and wear resistance both at room and high temperatures [2, 3].

However, the applications of these alloys have been limited due to their poor strength and wear resistance both at room and high temperatures [2, 3]. To overcome this problem, 316L alloys may be strengthened by a uniform dispersion of very hard fine ceramic particles [4]. The majority of reinforcing powders are directly added to metal matrix (i.e. ex-situ composites). The interface between the reinforcing particles and metal matrix is often a potential source of weakness due to differences in their thermal expansion coefficients. When ceramic particulates are incorporated by an ex-situ process, a thin oxide layer is formed on the surface of the particulates.

When ceramic particulates are incorporated by an ex-situ process, a thin oxide layer is formed on the surface of the particulates. This oxide layer affects the bonding with the metal matrix and may cause a crack that may propagate from the interface [5-8].To overcome such problems, several attempts have been made towards the development of in-situ MMCs in which the reinforcements are synthesized through a chemical reaction in the starting elemental materials [9]. In-situ formed ceramic reinforcements are thermodynamically stable in the matrix [10], leading to less degradation in elevated temperature applications. In particular, the reinforcement-matrix interface within in-situ MMCs are generally cleaner and more584compatible (i.e.

In particular, the reinforcement-matrix interface within in-situ MMCs are generally cleaner and more584compatible (i.e. free from gas absorption or oxidation) yielding stronger interfacial bonding and enhanced mechanical properties of the final products [11, 12].Apart from powder preparation, the use of additive manufacturing (AM) to consolidate metal powders into three-dimensional shapes with complex geometries has attracted considerable interest in the recent years. Selective Laser Melting (SLM) is an AM process that builds parts layer-by-layer with a moving laser beam that selectively fuses and consolidates thin layers of loose powder [13-15]. In previous work [16], we have studied the influence of starting TiC particle size and volume content in the developed 316L nanocomposite. The composites were developed by ex- situ powder preparation method.This study is aimed at the synthesis of in-situ TiC particles from pure Ti and pure C for reinforcement of 316L stainless steel.

The composites were developed by ex- situ powder preparation method.This study is aimed at the synthesis of in-situ TiC particles from pure Ti and pure C for reinforcement of 316L stainless steel. TiC particles can be formed by direct reaction between elemental Ti and graphite powder. Composite powders obtained after 35 h of milling were fabricated using SLM.Experimental proceduresFeedstock powder, and SLM processesThe starting powder materials were: water-atomized 316L having an irregular shape and an average particle size of 44 µm (Fig. 1a), Ti possessing a polygonal structure and a mean particle size of 10 µm (Fig. 1b), and pure graphite with a mean particle size of 2-12 µm (Fig. 1c). A powder mixture containing 90 vol.% 316L, 8 vol.% Ti, and 2 vol.% graphite was mechanically alloyed under a vacuum atmosphere using a Pulverisette 6 planetary high-energy ball mill.

1c). A powder mixture containing 90 vol.% 316L, 8 vol.% Ti, and 2 vol.% graphite was mechanically alloyed under a vacuum atmosphere using a Pulverisette 6 planetary high-energy ball mill. The milling conditions and parameters used for powder preparation are highlighted in Table 1.Specimens after 35 h of milling were used as a feedstock to be processed using the SLM parameters given in Table 2.

Cylindrical specimens with dimensions of 8 mm x 10 mm were produced.abcFigure 1: SEM micrograph of (a) 316L stainless steel powder; (b) Ti powder; (c) graphite powder.585Table 1: Milling conditions.AtmosphereRotational speed BallsBall/weight ratio Vesselvacuum82 RPM¼ stainless steel9:14.8" OD x 10" LongTable 2: SLM processing parameters.Power for area volume 100 WAtmosphere Applied energy density (J/mm3) 67, 100, 200, 300ArgonLayer thickness 50 µmMicrostructural characterizationsSamples were mechanically grounded and polished with progressively finer grit papers using the standard metallographic techniques and etched in Marble's Reagent for 10s. The microstructures of the composite powder and the SLM-processed samples were observed using a Nova 230 Nano scanning electron microscope (SEM). The chemical compositions of the powder were determined with an EDAX energy-dispersive X-ray spectroscopy (EDX) system.

The chemical compositions of the powder were determined with an EDAX energy-dispersive X-ray spectroscopy (EDX) system. The samples’ density was measured using the conventional Archimedes' principle.Mechanical testingVickers indentations was performed (Leco, LM800AT) at a load of 200 g at 15 random locations. Dry sliding wear tests were performed on the SLM-processed samples using a T50 ball-on-disk tribometer at room temperature as per the ASTM G99 standard. A 3-mm-diameter 52100 bearing steel ball was used as the counterface material. A load of 3 N was used during the tests. The friction unit was rotated at 840 rpm for 20 min; the rotation radius was 2 mm. The volume of material lost was measured using a ST400 white-light profilometer. The profilometer had a vertical resolution and accuracy of 8 and 80 nm, respectively, and a lateral resolution of 2 µm.

The profilometer had a vertical resolution and accuracy of 8 and 80 nm, respectively, and a lateral resolution of 2 µm. The cross-sectional area of the wear track was multiplied by the circumference of the wear track to get the total volume of material removed. This volume was then divided by the load applied and the total distance traveled during the test to calculate the wear rate of the material (mm3/N m). When the same test parameters are used for a number of samples, the wear rate can be used as a quantitative metric for comparing the wear resistances.ResultsPowder characterizationsFigure 2 shows the evolution of the morphologies of milled powders with the applied milling times. These micrographs show that the microstructural characteristics of milled powders (e.g., particle shape, particle size and its distribution, powder dispersion state) were significantly influenced by the milling times applied.

The soft 316L stainless steel particles were flattened in the early stage of milling and turned into flake-like particles after longer hours of milling. The Ti particles were able to undergo plastic deformation but the brittle C particles did not undergo plastic deformation or cold welding in the early stage of milling. The impact forces on the586particles imposed by the grinding medium caused severe plastic deformation of 316L particles and cold-welding to form larger particles. Further milling caused the particles to be work- hardened and the fracture mechanism to prevail over the cold welding process.2h6h20 h34hFigure 2: SEM micrograph of milled feedstock powder at different milling times.To further study the chemical composition of the milled powders, EDX characterization in selected areas of SEM images, was performed as indicated in Fig. 3. For a relatively short milling time of 2 h, the weight content of Ti and C detected on the particle surface was quite low.

3. For a relatively short milling time of 2 h, the weight content of Ti and C detected on the particle surface was quite low. This may have been due to low dissolution of 316L atoms in the Ti lattice in the initial stages of milling. The elemental concentration of Ti and C on the particle surface increased abruptly when the milling time increased. From this observation, we can conclude that a large number of regions where C atoms are dissolved in Ti lattice start to appear on the surfaces of milled powders, creating primary TiC solid solution zones. The EDX results showed that for milling up to 34 h however the weight fraction of 316L on the surface of milled particles decreased. This can be explained by considering that more 316L atoms penetrated deep from the surface into the interior regions of Ti particles and dissolved in the whole Ti volume, forming a TiC solid solution.

This could also be due to the fracture of 316L matrix powder particles causing fresh surfaces to be continuously created.2h6h20h34hFigure 3: EDX analysis of milled feedstock powder at different milling times.587SLM-processed composites characterizationsFigure 4 shows the low-magnification images for the etched composites samples using applied energy density of 67 (Fig. 4a) and 100 J/mm3 (Fig. 4b). Table 3 illustrates the density of the fabricated nanocomposite as a function of milling time. Clearly, the applied energy density affects the densification of the SLM-processed parts. On increasing the applied energy density from 67 to 100 J/mm3, the densification level increased to 98.62%, which was further dropped sharply to 88.81% at a higher energy density of 300 J/mm3. The corresponding microstructure at low energy densities (Fig. 4) showed that a number of balls were formed.

The corresponding microstructure at low energy densities (Fig. 4) showed that a number of balls were formed. Further details of SLM densification will be given in the discussion section.abFigure 4: Etched microstructure of the SLM-fabricated composites using different applied energy density (E): (a) 67; (b) 100.Table 3: Variation of relative density as a function of the applied energy density.SLM applied energy density (J/mm3)67100200300Relative density (%)95.5298.6289.5588.81Figure 5 shows the influence of the applied laser energy density on the characteristics microstructures in the SLM-processed composites. It was seen that grains are slightly of different sizes, may be due to the varying thermal heat fluxes during every scanning tracks. Nevertheless, a careful comparison of the microscopic features of these images revealed some distinct variations with the applied laser energy density.

Nevertheless, a careful comparison of the microscopic features of these images revealed some distinct variations with the applied laser energy density. With increasing the energy density, the grains tend to coarsen as a result of the lower scanning speed that favor slower solidification rate.588abcdFigure 5: Etched microstructure of the SLM-fabricated composites using different applied energy density (E): (a) 67; (b) 100; (c) 200; (d) 300.Mechanical behaviorTable 4 demonstrates the effect of the applied energy density of the 316L nanocomposites on the average microhardness and wear rate. The hardness were increased at 100 J/mm3 and then dropped at the applied energy densities of 200-300 J/mm3. The drop in hardness after applying higher energy density is a result of grain growth. Interestingly, the hardness exhibited similar trend as the relative density with respect to the applied energy density.

Interestingly, the hardness exhibited similar trend as the relative density with respect to the applied energy density. This suggests that the the relative density, applied energy density played an important role in influencing microhardness and wear performance of the nanocomposite.Table 4: Variation of microhardness and wear rate as a function of the applied energy density.SLM applied energy density (J/mm3)67100200300Microhardness (HV0.2) Wear rate (mm3/Nm)379.38±26.462.271385.94±36.462.061341.62±56.621.405370.83±51.731.596DiscussionIt was deduced that during laser processing, there is a strong likelihood of occurrence of the “balling phenomena” and unmelted particles (Fig. 4). This is the result of laser instability and insufficient applied energy density respectively [17]. The carbon content, used in this study, plays a critical role in determining processability of SLM.

This is the result of laser instability and insufficient applied energy density respectively [17]. The carbon content, used in this study, plays a critical role in determining processability of SLM. Indeed, the high carbon content also has adverse effect on SLM produced steel, since they portray a limited densification response. Wright et al. [18], found out that with increase in carbon contents, there is an increase in the thickness of the carbon layer accumulated on the melt surface. Just like the oxide layer, the carbon layer has a critical influence mostly on diminishing wettability and causing the melt to spheroidise rather than flow across the underlying surface. In addition to this, the brittleness of SLM manufactured high carbon content steel is increased by the formation of complex interfacial carbides at grain boundaries [19]. Childs et al. [20] concluded that the dissolution of carbides is favored by increasing the heat flow in the treated powder.

Childs et al. [20] concluded that the dissolution of carbides is favored by increasing the heat flow in the treated powder. On the same breath, increasing heat flow homogenizes distribution of alloying elements. Thus, the optimization of laser parameters is necessary. Due to this reason, high volumetric energy density for both elemental homogeneity and powder consolidation is achieved [15, 21, 22].589During the laser melting process, the metal powder absorbs the laser energy, raises its temperature, melt, and then solidifies when laser moves away, the whole process is less than one second. The heat also transfers away from the melt zone quickly. This causes the melt to cool down rapidly leading to short arrangement time and consequently porosity.

The heat also transfers away from the melt zone quickly. This causes the melt to cool down rapidly leading to short arrangement time and consequently porosity. In addition, because of the high cooling rate, the gas adsorbed on the surface of the powder, as well generated because of the reaction during the laser melting, have no time to escape from the molten pool and hence porosity was formed.When free interfaces exist in the system (e.g. liquid-liquid, liquid-gas interfaces) the temperature and compositional gradient will impose a surface tension gradient along the interface, which exerts a shear stress on the fluid. This induces a flow toward regions with higher values of γ termed Marangoni convection [23]. It is observed that during scanning “balling” has occurred. This is caused by surface tension driven by Marangoni convection (i.e. melting instability). Because of the strongly concentrated energy source, a large temperature gradient is often induced on the surface.

melting instability). Because of the strongly concentrated energy source, a large temperature gradient is often induced on the surface. If the surface tension γ depends on temperature, then the temperature gradient induces a surface flow that can have a considerable influence on the melting pattern. If the surface tension increases with temperature (dγ/ dT > 0), the liquid moves toward the energy beam on the surface, inducing narrower and deeper pool. However, if the surface tension decreases with temperature (dγ/ dT < 0), the surface flow moves away from the energy beam, leading to a wider and shallower pool.ConclusionsIn-situ TiC particle reinforced 316L stainless steel powder has been successfully used in the SLM process. The effects of SLM applied energy density on the microstructure and mechanical properties were studied. The SLM processed parts show a uniform fine equiaxed microstructure. Further optimizations of the processing parameters are necessary for future work.

The SLM processed parts show a uniform fine equiaxed microstructure. Further optimizations of the processing parameters are necessary for future work. We believe that in-situ MMCs fabricated by AM could be used to produce cheaper components that fulfill the increasing demand for high performance with their superior properties.AcknowledgmentOne of the authors, Bandar AlMangour, would like to extend his appreciation to the Saudi Basic Industries Corporation, which generously awarded financial support to him.References[1] AL-Mangour B. Powder Metallurgy of Stainless Steel: State-of-the Art, Challenges, and Development. Stainless Steel: Microstructure, Mechanical Properties and Methods of Application: Nova Science Publishers; 2015. p. pp. 37-80. [2] Krishna DSR, Sun Y. Effect of thermal oxidation conditions on tribological behaviour of titanium films on 316L stainless steel. Surface and Coatings Technology 2005;198:447-53. [3] Majumdar JD, Kumar A, Li L.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIn-Situ Veriﬁcation of 3D-Printed Electronics Using Deep Convolutional Neural NetworksDaniel Ahlers Department of Informatics, Group TAMS, University of Hamburg, Germany ahlers@informatik.uni-hamburg.de *AbstractPrinted electronics processes are becoming more stable and evolve into ﬁrst industrial applica- tions. These industrial applications require proper quality assurance to get a mostly autonomous production process. In this work, we present a new approach to inspect printed electronics and ensure their quality. Our hardware setup extends a fused ﬁlament fabrication (FFF) printer with an extruder for direct dispensing of conductive paste, a pick and place unit, and two cameras. The cameras take multiple images during printing.

The cameras take multiple images during printing. A trained neural network analyzes these pictures to separate the electronic wires from the plastic background. All separated images of a layer are combined to get a full view of the layer. Our algorithms then examine the detected wires to identify printing ﬂaws. The algorithms currently detect connection breaks, shorts, ﬁnd points that have not been reached, and evaluate the width of the printed wires.Keywords: printed electronics, visual inspection, neural network, fused ﬁlament fabrication1. IntroductionThe global additive manufacturing sector is constantly growing. In 2020, the market revenue was 12.8 billion US-Dollar, and it was growing signiﬁcantly even during the pandemic [1]. While other processes are already used on an industrial scale, 3D printed electronics are still evolving to be reliable enough for industrial applications.

While other processes are already used on an industrial scale, 3D printed electronics are still evolving to be reliable enough for industrial applications. 3D printed electronics is a specialized process to embed electronic circuits and components directly into printed objects. This technique can be used to print out prototypes or novel objects with integrated components like sensors. Although this process will never replace classical printed circuit boards (PCB), there are multiple use cases where 3D printed electronics adds additional freedom to product design. This freedom is also an issue for quality assurance since integrated wires are inaccessible after creating the object. Shorts and connection breaks caused by small imprecisions during printing are not detectable anymore. So, a proper in-situ quality assurance process is needed to use the 3D printed electronics process in industrial applications.

So, a proper in-situ quality assurance process is needed to use the 3D printed electronics process in industrial applications. This paper presents a new approach to verify 3D printed electronics during printing. Our FFF electronics printer takes multiple images during the printing process. A deep convolutional neural network then separates the conductive wires from the plastic substrate and the background. TheThis work was supported by the Federal Ministry for Economic Affairs and Energy in the ZIM-KamEl project.308algorithms then take the printed G-code and analyze the wires found by the network. The algo- rithms currently ﬁnd connection breaks, shorts, points that have not been reached by the printed wire, and evaluate each wire’s width. The rest of the paper is structured as follows. First, Sec. 2 summarizes the related work. Sec. 3 shortly describes the used hardware. Sec.

The rest of the paper is structured as follows. First, Sec. 2 summarizes the related work. Sec. 3 shortly describes the used hardware. Sec. 4 explains how the conductive wires are separated from the plastic substrate using a deep neural network. Sec. 5 introduces the algorithms used to ﬁnd the different issues in the printed wires. Sec. 6 shows some results and compares the outcome with previous works. The paper concludes with a summary and an outlook in Sec. 7.2. Related WorkThere are multiple techniques to integrate electronic wires into printed objects. With direct writing, a conductive silver paste is applied directly [2, 3]. Another way is directly embedding copper wires into the surface of an object [4]. Conductive wires can also be applied by aerosol jetting [5] or by ink-jetting [6]. The Stereolithography (SLA) process is also capable of printing 3D electronics [7]. Image processing is already broadly used for PCB inspection [8].

The Stereolithography (SLA) process is also capable of printing 3D electronics [7]. Image processing is already broadly used for PCB inspection [8]. Multiple works show the detec- tion of defect lines both on the surface [9] and on inner lines [10]. Faulty drilled holes are identiﬁed [9], and the correct placement of components is inspected [11]. In additive manufacturing, visual inspection is often used to improve the plastic/metal processes. Different approaches aim to quantify the geometric accuracy using different types of machine learning [12, 13, 14], inspect individual layers with a neural network [15], or monitor the printing process itself to detect defects [16, 17, 18]. Oleff et al. [19] presented an overview of the different approaches for process monitoring in additive manufacturing using different sensors like cameras, thermal cameras, accelerometers, 3D cameras, microphones, and other sensors. For conductive wires, Salary et al.

For conductive wires, Salary et al. [20] showed a method to quantify the attributes in aerosol jetting. This includes the line width, line density, line edge quality/smoothness, overspray, line discontinuity, and internal connectivity. They also showed that the estimated resistance could be calculated using shape-from- shading [21]. This work extends our previous work for optical veriﬁcation of 3D printed electronics [22]. A support vector machine (SVM) was trained on every layer to classify the pixels into conductive or not conductive pixels by only looking at their color value. This classiﬁcation is a bit noisy and has problems with dark regions near the conductive material and reﬂections on top of the wires. A sliding window is shifted along the planned extrusion, and all conductive pixels are summed up to ﬁnd connection breaks. If the number of pixels falls below a threshold, the section is classiﬁed as a connection break, nevertheless if there is a broken connection.3.

If the number of pixels falls below a threshold, the section is classiﬁed as a connection break, nevertheless if there is a broken connection.3. Hardware SetupOur new hardware setup to fully automatically print 3D electronics is a modiﬁed E3D ToolChanger [23]. The open-source ToolChanger is a platform that can attach four different tools to a motion system. These tools can be parked and picked up again. Fig. 1 shows our printer with the different tools used for conductive printing. All tools only move along the X-, Y-plane while the printbed moves along the Z-axis. A camera is attached to the toolhead facing downwards to move around with the attached tools permanently. Currently, two FFF extruder printheads are attached to the printer. One is directly driven, and the other is a bowden driven extruder.

Currently, two FFF extruder printheads are attached to the printer. One is directly driven, and the other is a bowden driven extruder. A pneumatic conduc-309Figure 1 – The modiﬁed E3D ToolChanger with the FFF printhead, the vacuum pickup tool, the rotating conductive extruder, the camera attached to the toolhead, and the component tray with the second camera next to it.tive ink dispenser can extrude conductive silver paste to create electronic wires when pressure is applied. A vacuum pick and place tool is used to grab components from the component tray and place them into the objects. The component tray can be moved up and down to pick up the components. The second camera is used to align components and calibrate the printer’s different tools, similar to our previous work [24]. Since all tools have a different length, the camera can be moved up and down to shift the focus to the tool’s tip.

Since all tools have a different length, the camera can be moved up and down to shift the focus to the tool’s tip. Both cameras are a Raspberry Pi Camera Module V2 with a resolution of 3280 × 2464 pixels, a short focus distance, and an attached led ring to have reproducible light conditions. Since these cheap cameras tend to produce distorted images, we calibrated the intrinsic parameters of the cameras and corrected every image before further processing.4. Wire SegmentationThe images for the veriﬁcation of the 3D printed electronics are taken during printing each time a layer is ﬁnished. The camera in the print head takes multiple images per layer to cover the whole printed object. The pictures are aligned and recorded automatically by our software [22]. For better matching, the tiles are taken with an overlap to align them automatically. This is necessary because of positioning inaccuracies from the belt-driven gantry system.

For better matching, the tiles are taken with an overlap to align them automatically. This is necessary because of positioning inaccuracies from the belt-driven gantry system. The tiles are combined into a large image where each pixel covers 20 µm × 20 µm. The images are then fed into the neural network to separate the conductive wires from the plastic substrate and the background. Such a large image can not be feed directly into the network, and scaling it down to a usable size would lead to a loss of detail. So the image of the whole layer is310128166432642562561286432646416Bottleneck128256321285126412832Softmax5121625651212832256Figure 2 – The structure of the U-net with a 512 × 512 pixel input and output. All layers use ReLu as an activation function except the output layer, which uses Sigmoid.again split up into multiple tiles where each tile is 512 × 512 pixels.

All layers use ReLu as an activation function except the output layer, which uses Sigmoid.again split up into multiple tiles where each tile is 512 × 512 pixels. Because convolution layers with active padding add zeros on the edges to keep the input and output size the same, the detection from the network is wrong along the edges of the tiles. This is compensated by adding an overlap of 32 pixels to the tiles. The overlap is removed when rearranging the output tiles back into one image after all tiles are processed. Different network architectures were implemented and tested with the same dataset to ﬁnd the best working network architecture. The tree architectures are U-net [25], DenseNet56 [26], and DeepLabv3+ [27], all implemented with TensorFlow. While all architectures performed similarly in detection rates, the U-net architecture was chosen because it performed best on small detail regions like plastic strings and small gaps in the material.

The U-net takes an image with the tree color channels as an input. The contraction part contains four blocks, each with two convolution layers and a max-pooling layer. The bottleneck only contains two convolution layers. The layers are then expanded with four blocks to get back to the input resolution. Each block starts with a transposed convolution, followed by the concatenation with the bypass and two convolution layers. All convolution layers use Relu as an activation function. Lastly, a convolution layer outputs three probabilities for each pixel using a Sigmoid activation function. The three classes are conductive wire, plastic substrate, or background. Fig. 2 illustrates the layout of the network as described previously. The training data for the network is generated from images recorded for documentation purposes in previous prints. The images are extracted from six different prints. They are printed in white and black plastic.

The images are extracted from six different prints. They are printed in white and black plastic. The image tagging of the electronic wires, the plastic, and the background was done by hand. To get the images in the same size as the network, they are split up into 512 × 512 pixel sub-images with an overlap of 256 pixels. Since the dataset has only two colors, a copy of the images is added where the color of the plastic was changed with GIMP. Printing more colors with the printer instead of artiﬁcially creating images was planned but could not be realized due to a closed lab during the pandemic. More data augmentation is added by rotating a small proportion3110.2Training loss Validation loss) E C C S (0.1s s o L0246810121416 Epochs182022242628Figure 3 – The plotted loss of the training and validation sets during the epochs.of the images, changing their brightness and contrast.

The resulting dataset has roughly 20,000 images split into 80% training set and 20% validation set. The network was trained for 30 epochs and converged pretty fast. The resulting training and validation loss from the Sparse Categorical Cross Entropy (SCCE) function is plotted in Fig 3.5. Wire VeriﬁcationWith the separated image from the network, the veriﬁcation algorithms can be applied. First, the printed wires are extracted from the G-code ﬁle. Then they are converted into the image coordinate system. This is possible because the positions where the corrected pictures are taken and the pixels per mm are known. Fig. 4 shows the output from the neural network overlayed with the G-code instructions. Since the wires in the G-code have no information about the relationship of the different lines, they are grouped before starting the veriﬁcation. The grouping algorithm tests every line if one endpoint is near any other endpoint of another line.

The grouping algorithm tests every line if one endpoint is near any other endpoint of another line. If two points are closer together than a threshold (1/2 wire width), they belong to the same wire and are grouped. If a line has points close to two groups, they are merged because they belong to the same wire. When all lines are grouped, the printed wires can be evaluated. The following paragraphs describe how the different appearing problems can be identiﬁed. All tests are repeated for all grouped wires that were extracted from the G-code.Shorts are unwanted connections between two or more wires often caused by too much con- ductive material, imprecise movement, or a leaking conductive extruder. A short can be found by checking if any other wire can be reached from the current wire. To test this, the segmented image from the network is taken and the starting point of the wire is ﬂood-ﬁlled.

A short can be found by checking if any other wire can be reached from the current wire. To test this, the segmented image from the network is taken and the starting point of the wire is ﬂood-ﬁlled. Flood-ﬁll recursively marks every neighbor pixel that is classiﬁed the same as the starting pixel. Then, every point of all other wires is checked if they got ﬁlled. If they got ﬁlled, these two wires are connected and have a short.31230Figure 4 – Left: The output of the wire segmentation where gray areas are identiﬁed as wires and black areas are plastic substrate. Right: The overlayed wires extracted from the G-code in red.Connection breaks are unwanted interruptions in the conductive material. Unconstant extru- sions or small plastic strings often cause these issues. Connection breaks are also checked with the ﬂood-ﬁll technique.

Unconstant extru- sions or small plastic strings often cause these issues. Connection breaks are also checked with the ﬂood-ﬁll technique. The starting point of the wire is ﬂood-ﬁlled and then every point of this wire that got extracted from the G-code is checked if they got ﬁlled correctly. If a point got not ﬁlled, there is no connection between the starting point and this point, so there must be a connection break somewhere in between. The part of the wire that is broken can be found by looking for the line extracted from the G-code where one end of this line is ﬁlled and the other is not.Points that are not reached often appear on sharp corners when the material is slightly off due to an imprecise motion system or on the ends of a wire due to inconsistent material ﬂow. They can be found by checking if all points from the G-code of a wire are classiﬁed as conductive material. If not, this point is not reached during printing.

They can be found by checking if all points from the G-code of a wire are classiﬁed as conductive material. If not, this point is not reached during printing. While a connection break or a short is a clear fault, a point that was not reached can still be tolerable since there is a general connection in the wire, and reaching this exact point might not be important.Wire width is often crucial since the width of the wire inﬂuences its resistance. A wire can be connected but have a thin section along its path. To measure a single wire and not accidentally include another one nearby, all other wires are removed from a copy of the segmented image. Then each wire segment from the G-code is extracted into a single image and rotated vertically. So the image only contains a segment from one corner to another. The width is limited to four times the conﬁgured extrusion width of the conductive wire to reduce the impact of the corners.

So the image only contains a segment from one corner to another. The width is limited to four times the conﬁgured extrusion width of the conductive wire to reduce the impact of the corners. Then all pixels are then summed up in every pixel row to get the width of the wire in this pixel row. The pixel values are then converted into mm to get real-world values. Fig. 5 illustrates the evaluation of one wire segment.313Figure 5 – Left: The isolated wire group. Middle: The cropped and rotated wire segment. Right: A plot of the summed-up pixel rows representing the wire segments width (mm) in printing direction.6. EvaluationThe wire segmentation is evaluated by comparing the detection performance of the neural network with our previous work [22]. Multiple images are processed by the neural network and the SVM and then compared with a manually labeled version of the images.

Multiple images are processed by the neural network and the SVM and then compared with a manually labeled version of the images. Table 1 lists the correct classi- ﬁed pixels, the false-negative pixels, and the false-positive pixels for both approaches. Note that the summed-up pixels are only classiﬁed as conductive wire by the different approaches and the per- centage rate is compared to the actual number of existing conductive pixels in the manually labeled images. False-positive means that pixels are marked as wires but are no wires. False-negative is the other way round where wire pixels are marked as plastic or background. Both cases are a prob- lem since false positive pixels tend to indicate a short when there are none and can lead to missed connection breaks, while false-negative pixels might lead to missed shorts and detect connection breaks that are nonexistent.

The results show that the rate of correctly classiﬁed and false-negative pixels stays similar to the SVM approach while the false-positive rate drops signiﬁcantly. When inspecting which regions are wrongly classiﬁed, it shows that dark regions are generally less often wrongly classiﬁed. However, when the dark region lies between two conductive wires, the classi- ﬁcation sometimes fails and marks the region between these two wires also as wire. Reﬂections are less problematic, but very bright ones are still sometimes wrongly classiﬁed.Table 1 – Pixel classiﬁcation of conductive wires in absolute pixels and the detection rate.correct U-net 916304 (96.6%) SVM 904626 (95.4%)false-positive 60008 (6.3%) 279645 (29.5%)false-negative 32407 (3.4%) 44085 (4.6%)314Figure 6 – Four different defects and the output from the veriﬁcation algorithms.

From top left to bottom right: A connection break created by a plastic string over the conductive material, a connection break from missing material, a short from an accidentally printed line, and a short from a misplaced conductive line. In the results, connection breaks are marked red, shorted wires are drawn yellow, and points that are not reached by the conductive material are marked orange.To evaluate the veriﬁcation algorithms, we created images with intentional defects that these algo- rithms should ﬁnd. The defects are connection breaks due to inconsistent extrusions, connection breaks produced by thin plastic strings from leaking plastic extruders, and different shorts. Since we had no pictures of shorted connections, we created shorts using GIMP to test our algorithms. Fig. 6 shows different defects with the output of the wire veriﬁcation algorithms. It can be seen that our algorithms detect connection breaks correctly and marks the broken wire segments in red.

Fig. 6 shows different defects with the output of the wire veriﬁcation algorithms. It can be seen that our algorithms detect connection breaks correctly and marks the broken wire segments in red. The shorts from the accidentally printed line and the misplaced conductive line are also identi- ﬁed and marked in yellow to indicate that these two wires are shorted. The algorithms also found three points that are not reached and marked them with orange circles. While the point on the ﬁrst short is not reached because the material is not printed until the end, the other two are not reached because the conductive material is not in its intended position. The extra time that it takes to run the whole veriﬁcation process is crucial. The wire segmentation with the neural network takes ∼6 seconds, while the wire veriﬁcation takes ∼1 second for one layer. Both measured on a notebook with a 4 core 3.6 GHz CPU.

The wire segmentation with the neural network takes ∼6 seconds, while the wire veriﬁcation takes ∼1 second for one layer. Both measured on a notebook with a 4 core 3.6 GHz CPU. So the inﬂuence on the print time per layer is less than 1% which is similar to our previous work and does not increase the printing time signiﬁcantly. However, capturing the images during the printing still takes about 10% extra time on every layer, increasing the printing time signiﬁcantly.3157. Conclusion and OutlookIn this work, we showed our new approach to verify 3D printed electronic wires while printing it. The approach focuses on FFF printed plastic objects with pressure extruded conductive silver ink but could also be applied to different printing techniques, substrates, and wire embedding techniques. It compares the G-code instructions with the actual printed object without any needed supervision and stores all data for a high-resolution documentation.

It compares the G-code instructions with the actual printed object without any needed supervision and stores all data for a high-resolution documentation. A professional can look at the found errors and decide if the device can be used or not. Knowing the wire width of each wire can be used to improve the printing process. Our future work will focus on improving the wire veriﬁcation to identify the exact position of a short or connection break. A found connection break could be printed again to get a stable connec- tion during printing and close the control loop. Since shorts and connection breaks can also appear between layers, we want to connect these individual layers to ﬁnd defects that appear between two adjacent layers. The correct placement of SMD components also needs further inspection. Here, the printed footprints and the placement of the components need to be veriﬁed.

The correct placement of SMD components also needs further inspection. Here, the printed footprints and the placement of the components need to be veriﬁed. The wire thickness measurement can also be fed back into the printer to adjust the material ﬂow during printing.References[1] Terry Wohlers, Robert Ian Campbell, Olaf Diegel, Ray Huff, and Joseph Kowen. WohlersReport 2021. Wohlers Associates, Fort Collins, Colorado, 2021.[2] Cassie Gutierrez, Rudy Salas, Gustavo Hernandez, Dan Muse, Richard Olivas, Eric Mac- Donald, Michale D Irwin, Ryan Wicker, K Churck, M Newton, and Brian Zufelt. CubeSat Fabrication through Additive Manufacturing and Micro-Dispensing. Proceedings from the IMAPS Symposium, 2011.[3] Florens Wasserfall, Daniel Ahlers, Norman Hendrich, and Jianwei Zhang. 3D-Printable Elec- tronics - Integration of SMD Placement and Wiring into the Slicing Process for FDM Fab- rication.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIncorporation of Automated Ball Indentation Methodology for Studying Powder Bed Fabricated 304L Stainless SteelSreekar Karnati, Jack Hoerchler, Aaron Flood, Frank LiouDepartment of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409AbstractAutomated Ball Indentation (ABI) is a viable method for estimating the ductility, yield stress, and ultimate stress, among other metrics, in different metallic materials. Currently, ABI data analysis utilizes Holloman’s Power Law to model the plastic region of the true stress-true strain curve. While this formulation is accurate for some materials, its relevance for additively manufactured austenitic stainless steels, such as 304L, needed investigation.

While this formulation is accurate for some materials, its relevance for additively manufactured austenitic stainless steels, such as 304L, needed investigation. The deviation of the material’s plastic behavior from the Power Law was investigated. In order to better model this behavior, both the Voce and Ludwigson formulation were investigated. These formulations were tested for both wrought and additively manufactured 304L stainless steel. Regression analysis was used to choose the appropriate fit. The chosen formulation was then used to generate a material model to simulate the ABI process. These simulations were validated through experimental analysis.IntroductionTraditionally, metal components for various engineering applications are fabricated from a raw stock material, through steps of machining and assembly. The choice of raw material usually depends on the requirements of the engineering application.

The choice of raw material usually depends on the requirements of the engineering application. The properties of these raw materials are established through extensive characterization. Post-fabrication, the same stock properties are expected from the machined and assembled components. This approach has been successful and is widely used to date. However, the buy to fly ratios for fabricating components in this manner can be high. Especially, fabricating components from tough and expensive materials can be challenging. As a solution, Additive manufacturing has been identified as a potential means to fabricate such components in an economical fashion.infeasible complex geometries from multiple varieties of material. It is also fundamentally different from the traditional approaches. Instead of removing material from a raw stock, AM lets one fabricate individual components piece by piece by adding material in a layer by layer fashion.

Instead of removing material from a raw stock, AM lets one fabricate individual components piece by piece by adding material in a layer by layer fashion. Unlike conventionally used subtractive manufacturing, the properties of the AM parts cannot be traced back to stock material. Additionally, AM materials/parts have been known to possess excellent yet anisotropic properties. The properties of AM material have been known to vary with attributes such as part size, build strategy, laser power, scan speed etc. [1–6].

The properties of AM material have been known to vary with attributes such as part size, build strategy, laser power, scan speed etc. [1–6]. The lack of the possibility to trace back the properties to a stock material and the wide variation within the material makes part inspection and qualification quite challenging.Additive Manufacturing (AM)is capable of fabricatingtraditionally©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1349Destructive testing of AM components to assess the influence and variation caused by the various attributes of the fabrication process can be cost prohibitive, counterproductive and difficult. In order to economically investigate the material performance of AM material, miniature tensile testing was utilized [7–9].

In order to economically investigate the material performance of AM material, miniature tensile testing was utilized [7–9]. While miniature tensile testing was successful in characterizing AM material, it is still a destructive mode of testing. The relevance of these results may also be questioned as it is suspected that the part geometry has an influence on part properties. If the geometry from which the mini-tensile specimens were sampled, happens to be substantially different from the part under scrutiny, the data from mini-tensile testing might be invalid. For reasons of such ambiguity, there is a need for non-destructive means to characterize and inspect AM material and components.Currently, there exists a wide range of capabilities for performing a non-destructive evaluation for inspection and characterization. Techniques involving, acoustic sensors, infrared sensors, X-rays, microwaves, eddy currents, natural frequency etc.

Techniques involving, acoustic sensors, infrared sensors, X-rays, microwaves, eddy currents, natural frequency etc. are used from routine inspection to characterization[1,10–15]. Automated Ball Indentation (ABI) is one of such non- destructive testing methodologies. It is capable of estimating multiple material properties. ABI techniques utilize ball indentation data and empirical analysis to estimate various mechanical properties. The ABI technique was developed by Haggag et.al. and is currently standardized through ASTM[16]. The technique has also been proven to be effective for fracture toughness estimation[17].Multiple researchers have demonstrated the viability of utilizing the ABI technique to estimate the mechanical properties of metals. It was claimed to be capable of monitoring the structural integrity of components in nuclear, aerospace, chemical and process plants. This was demonstrated by investigating the viability on materials favorite to these industries[18].

This was demonstrated by investigating the viability on materials favorite to these industries[18]. The technique was also found to be capable of identifying spatial gradients in material properties from welding[19]. His technique was also capable of identifying the influence of aging time on mechanical properties of Inconel 625 alloy[20]. While the technique shows great promise, the relevance and validity of this technique for every material requires investigation.In this paper, the validity of the ABI technique for investigating the mechanical properties of powder bed fabricated 304L stainless steel were investigated. The validity of some assumptions central to the empirical analyses involved in processing ABI data was also investigated. For a better understanding of the technique, the indentation process was also simulated using finite element analysis techniques.

For a better understanding of the technique, the indentation process was also simulated using finite element analysis techniques. Mechanical properties gathered from destructive testing were compared with values estimated through ABI testing.TheoryThe ABI technique uses load vs. depth of indentation data gathered from guided loading and unloading cycles of a ball indenter to estimate various material properties. A schematic layout of the empirical analysis involved in estimating yield strength, engineering ultimate strength, strength coefficient, strain hardening exponent and Meyer’s index is shown in Figure 1. The equations involved in the empirical analysis are shown in Figure 2.©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1350The depth of indentation data is used to estimate the chordal diameter of indentation.

Schematic layout of data processing involved in estimating yield strength, strength coefficient, strain hardening exponent and Meyer’s index[16]Figure 2. Representation of True (pileup, dt) and Chordal diameters (in-plane, dp) of an ABI impression[17].The Holloman’s power law (Equation 1) used to the model the plastic portion of the true stress-true strain curve is not an accurate formulation for austenitic stainless steels such as 304L stainless steel. A significant deviation from estimated stress value is noticed for low values of strain. The Ludwigson (Equation 1) and Voce formulations (Equation 3) are more accurate in estimating the full plastic behavior of 304L stainless. This was demonstrated at multiple©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1351temperatures for 316L austenitic stainless steel by Choudary et.al.[21].

The equations of the formulations are shown below. Where is the calculated true plastic strain, k, n, k1, n1, A and B are empirical parameters.

Equations involved in the ABI empirical analysis analyses, these equations are used in the sequence shown in Figure 1 [17].Experimental setupThe stainless steel analyzed in this paper was procured in three states, (1) powder bed fabricated in as-built state, (2) powder bed fabricated and heat treated, and (3) wrought, cold rolled and annealed. The stainless steel fabricated using the laser powder bed fusion process was©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1352produced on a Renishaw AM250 powder bed machine. The parameters used for fabrication were optimized for density and smooth downward facing surfaces. 304L stainless steel powder in the size range of 15-45 micron was used for fabrication. The heat treated material was produced by subjecting the as-built material to 300C for 1 hour.

304L stainless steel powder in the size range of 15-45 micron was used for fabrication. The heat treated material was produced by subjecting the as-built material to 300C for 1 hour. The wrought material was commercially procured.The ABI indentation was performed using a commercially procured machine. The machine was designed and produced by ABI Services LLC. The indentations were performed using a 0.75mm diameter WC ball indenter. A schematic representation of the loading and unloading cycles is shown in Figure 4. The current implementation of ABI testing involves eight cycles of loading and unloading. Each cycle has the same amount of depth of indentation. The depth of indentation for each cycle is the ratio of peak depth (input as a percentage of the radius of the indenter) and indentation and the total number of cycles (8 cycles). The unloading at the end of each cycle is intended to relieve any elastic deformation and accurately estimate plastic deformation.

The unloading at the end of each cycle is intended to relieve any elastic deformation and accurately estimate plastic deformation. The Figure 3 highlights the relationship of varying %radius and unloading percentage with load and depth of penetration data.Figure 4.An illustration of the load vs. depth of indentation (displacement) obtained using controlled loading and unloading of a ball indenter[17]A finite element model was created using ANSYS APDL 18.1 simulation software in order to simulate the plastic deformation occurred during the ball indentation process. Using ANSYS, a script has been written to simulate the ABI testing process. This included a ball indenter situated over a block to be indented on. The ball indenter has been meshed in order to define the contact surface, and the test block has been meshed to optimize accuracy and calculation time using mapped meshing techniques. The meshing schemes are shown together in Figure 5.

The meshing schemes are shown together in Figure 5. Contact and Target surfaces have been defined as the areas on the indenter and the top face of the base block respectively. These surfaces will allow for the indenter to deform the target surface on the base block. The ball was meshed using fine free meshing. This allows the program to create a non- uniform map of nodes across the indenter to evenly place nodes on the contact face. The base was meshed with a mapped cubic 8 node mesh. This allowed for a finer mesh to be created closest to ©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1353the point of indentation on the block. Doing this improves the quality of the results in the area around the indentation, as well as speeds up calculation time due to fewer elements being used in non-critical locations.

Doing this improves the quality of the results in the area around the indentation, as well as speeds up calculation time due to fewer elements being used in non-critical locations. A quadratic 20 node 3D element was used to simulate the material behavior during indentation.Figure 5. A meshed model of the ball indenter and the test block.The material parameters of the ball were the same as the values used by the ABI analysis software. The model currently accounts for the material properties of the test block by including modulus of elasticity of 304L (from literature), the nonlinear plastic deformation section of the stress-strain curve (experimental, gathered from testing 6mm X 1mm X 1mm Miniature tensile specimens specimen of SLM material [9]), and the Poisson’s ratio of 304L (from literature). The model is set to a constant temperature and temperature dependent properties are ignored due to it having a negligible effect on results.

The model is set to a constant temperature and temperature dependent properties are ignored due to it having a negligible effect on results. A transient analysis was used to replicate the indentation process. 200-time steps were used in order to displace the indenter in steps down into the block and back up to a position that is clear of the block. These steps were based on the experimental data. The location of the ball during the loading and unloading cycles was identified and used as the input for indentation simulation.The test block was created to be larger than the interaction volume for ball indentation tests in which the material is affected. This interaction volume was based on the cut-off that plastic deformation wouldn’t go beyond twice the diameter of the indenter. This allowed for directly fixed constraints to be used on the outer faces of the block as well as the bottom face of the block.

This allowed for directly fixed constraints to be used on the outer faces of the block as well as the bottom face of the block. The indenter is fixed above the block base at all of the individual nodes of the block. The indenter is then moved downward/upward in a cyclic fashion over the course of 200 individual steps, end of indentation the indenter moved upward to remove the load on the block. Each step resets the©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1354displacement of the indenter to a new location, moving the indenter. The indenter movement can be realized from the plot shown in Figure 6. This corresponds to an indentation rate of 0.005mm/s.0501001502000r e t n e d n0.020.04il la b f o n o i t a c o L0.060.080.10.12Step #Figure 6.

This corresponds to an indentation rate of 0.005mm/s.0501001502000r e t n e d n0.020.04il la b f o n o i t a c o L0.060.080.10.12Step #Figure 6. Location of the ball indenter during the course of a 200 step indentation simulation. The peak depth of indentation is 24% of the radius and the total number of loading-unloading cycles are 8.Results and discussionsYield strength estimationThe current section of analysis involved the evaluation of the empirical parameters needed to model the linear relationship between ‘A’ parameter and tensile yield strength. The validity of the evaluated model with changing indenter size and changing interaction volume was also investigated. The setup of the linear model between ‘A’ parameter and yield strength data was based on tensile data gathered from as-built AM material and heat treated AM material.

The setup of the linear model between ‘A’ parameter and yield strength data was based on tensile data gathered from as-built AM material and heat treated AM material. Data from 25 specimens in as-built condition [22] and 13 specimens in heat treated condition [23] was used for the setup of the linear model. A set of 3 indentations were performed on each these broken specimens. The ‘A’ parameter gathered from ABI testing and corresponding tensile yield strength measurements were linearly fit to evaluate the empirical constants of the linear relationship. The validity of the obtained fit was cross-checked by comparing indention and tensile data from a new set of 12 specimens in as-built condition.ABI literature states the material slope parameter for stainless steels is 0.189 [16], while the intercept value is 0. By forcing the intercept value to be zero, the material slope parameter was found to be 0.1936. However, this was not a good fit.

By forcing the intercept value to be zero, the material slope parameter was found to be 0.1936. However, this was not a good fit. A linear fit with non-zero intercept yielded a good fit. This model was successful in estimating the yield strength values of the second set of indentations. The raw data, the linear fits with zero and non-zero intercepts and their corresponding R2 values are shown in Figure 7. Although the tensile data spans over a wide range, two clusters©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1355250of data points can be seen in the plot. These clusters are representative of the state of the material that was considered in this study. The lower end is the heat treated material and the higher end is the as-built material.

These clusters are representative of the state of the material that was considered in this study. The lower end is the heat treated material and the higher end is the as-built material. The validity of the model for material with properties outside the clusters is yet to be investigated.Heat treated + As built550.0500.0y = 0.1465x + 105.52 R² = 0.9431S Y e450.0l i s n e T400.0y = 0.1936x R² = -0.734350.0300.01500.01700.01900.02100.02300.02500.02700.02900.0A parameterFigure 7. Linear fits between tensile yield strength and A parameter. The fit with non-zero intercept offers good correlation (red)Table 1. Yield strength measurements taken using tensile testing and their corresponding ABI estimates.

The fit with non-zero intercept offers good correlation (red)Table 1. Yield strength measurements taken using tensile testing and their corresponding ABI estimates. Mean and standard deviation in parenthesesYield Strength (MPa)Tensile measuredABI estimated500.4424;458;422;448 (438,15.4)519.5433;410;427;405 (418,11.5)503.8434;453;424;441 (438, 10.5)The linear model with the intercept was used as the model estimating yield strength. The new data set gathered to investigate the validity of the model is shown in Table 1. A significant difference can be seen between tensile measurements and ABI estimates. The ABI technique underpredicts the yield strength values. A significant scatter can also be noticed in the ABI estimates. The impact of varying indenter radius on ABI measurements and validity of the linear model at each size was investigated. By increasing the size of the indenter, decreasing values of ‘A’ parameter were obtained.

By increasing the size of the indenter, decreasing values of ‘A’ parameter were obtained. Consequently, lower values of yield strength estimates were©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1356obtained from ABI. The values of ‘A’ parameter and yield strength estimates from ABI linear models are shown in Figure 8. The models with zero and non-zero intercepts were used. A slope of 0.189 was used for the model with zero intercept. The model with non-zero intercept is shown in Figure 7.Figure 8. Decreasing values of ABI estimates with increasing indenter diameterFlow curve formulationThe flow curve modeling is an integral part of the empirical analysis performed by the ABI software. As a default, the Holloman power law is used to model the plastic section of the true stress-true strain curve.

As a default, the Holloman power law is used to model the plastic section of the true stress-true strain curve. However, the power law is not a good representation of the plastic flow of 304L stainless. The same can be seen from the Figure 9. The true stress-true strain data calculated from mini-tensile specimens of 3mm gage length and 1mmx1mm cross-section (both wrought and AM, [9]) were used to fit Ludwigson and Voce models to the plastic portion of the stress-strain curve. While Ludwigson (Equation 4, As-built AM) was better than Holloman fit, it was still not completely representative. Voce model (Equation 5, As-built AM) was found to be the best fit. The same results were seen for both wrought and AM materials.

Voce model (Equation 5, As-built AM) was found to be the best fit. The same results were seen for both wrought and AM materials. The goodness of it can be seen from the residual plots shown in Figure 10 and Figure 11.Equation 4𝜎𝜎 = 1411.1(𝜀𝜀𝑝𝑝)0.341+ 𝑒𝑒(5.7−18.2𝜀𝜀𝑝𝑝)Equation 5−0.964𝜀𝜀𝑝𝑝𝜎𝜎 = 2137.2 − (2137.2 − 586.9)𝑒𝑒©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1357Figure 9. Ln-ln plot of true stress-true strain curves and their Holloman power law fitsFigure 10. Residual plots of Ludwigson model for as-built AM material.

Ln-ln plot of true stress-true strain curves and their Holloman power law fitsFigure 10. Residual plots of Ludwigson model for as-built AM material. The pattern in versus fits and versus order graphs indicate un-modeled data in the residuals.©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1358Normal Probability PlotVersus Fits99.99995.0t n e c r e P905010l a u d i s e R2.50.02.515.00.0110505106007008009001000ResidualFitted ValueHistogramVersus Ordery c n e u q e r F806040l a u d i s e R5.02.50.02.5205.005.253.501.750.00Residual1.753.505.25120 040 060 080 0100 0120 0140 0160 0180 02 00 02 20 0Observation OrderFigure 11. Residual plots of Voce model for as-built AM material.

Residual plots of Voce model for as-built AM material. Indicative of a good fit of true stress by using Voce modelAlthough the Voce formulation is a more accurate fit for modeling the flow curve, it is to be noted that the deviation of the power law from experimental data is only for low values of strain (Figure 9). This means the yield strength estimation from the powder law is infeasible (stress at true plastic strain=0). However, the ultimate strength estimation might still be valid. The power law might be used for strain values higher than a threshold (0.135 mm/mm in this case). The ultimate strength values estimated using the power law are shown in Table 2. While there is still a difference between the measured and estimated values, the difference is low in comparison to yield strength estimation.Table 2.

While there is still a difference between the measured and estimated values, the difference is low in comparison to yield strength estimation.Table 2. Ultimate strength measurements taken using tensile testing, and their corresponding ABI estimatesUltimate Tensile Strength (MPa)Tensile measured ABI estimated (Mean, std. dev.)673.9709;708;694;695 (701, 7)684.6609;609;661;661 (635, 26)674.1655;656;695;695 ( 675, 19.7)©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1359Finite element method simulationCurrently, all of the material properties used in the simulation are time independent. The modulus of elasticity has been input as 200 GPa, and the Poisson’s ratio has been input as 0.33, both standard values for 304L stainless steel.

The modulus of elasticity has been input as 200 GPa, and the Poisson’s ratio has been input as 0.33, both standard values for 304L stainless steel. The exact values of modulus and Poisson ratio of SLM fabricated 304L stainless steel is expected to be different from those of wrought material. For better accuracy, these values are to be experimentally evaluated and inputted into the simulation. Plastic deformation was accounted for by inputting a table to calculate the nonlinear plastic deformation region of the true-stress vs. true-strain curve for 304L, shown in Table 3.The required number of elements was identified by running simulations with various mesh sizes. The load-displacement (depth of indentation) curve obtained at the end of simulation was used as the norm for identifying the required mesh size. Meshes with total nodes of 20k to 80k were used for simulations.

Meshes with total nodes of 20k to 80k were used for simulations. A snapshot of Von Mises stress and Von Mises strain in the test block (80k nodes) and indenter post indentation can be seen in Figure 12. The calculated load- displacement curves from the each of those simulations are shown in Figure 13. The difference between the load values at same positions of the indenter was used to identify convergence. The average and standard deviation of these differences are shown in Table 4. A mesh size of 60K nodes was expected to be sufficient to achieve convergence in estimating the load-depth of indentation plot.Table 3.

Contour plot of Von Mises stress and strain in the indenter and test block post indentation (mesh size of 80k nodes).FORCE V DISPLACEMENT60050020000MESH)N40030000MESH( d a o L30020040000MESH50000MESH60000MESH10070000MESH000.020.040.060.080.10.120.1480000MESHDepth of indentationFigure 13. Simulated load-depth of indentation curves obtained through mimicking ABI test on blocks of various mesh sizes.Table 4. Average and standard deviation of the difference in load values at same depths of indentation for cases shown in Figure 2.10. A mesh size of 60K nodes is expected to achieve convergence80k-20k 80k-30k 80k-40k 80k-50k 80k-60k 80k-70kMean5.932.942.442.391.631.56Std.

A mesh size of 60K nodes is expected to achieve convergence80k-20k 80k-30k 80k-40k 80k-50k 80k-60k 80k-70kMean5.932.942.442.391.631.56Std. Dev5.893.002.182.071.221.37©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1361The chordal diameter of the indentation is calculated through an analytical model (see Figure 2 and Figure 3). The estimated value is used for calculating flow stress and plastic strain values. The validity of this estimation is difficult to evaluate as an accurate measurement of chordal/ indentation diameter can prove to be difficult. In order to assess the validity through simulation, convergence needs to be ensured for plastic flow as well. The values of these diameters were calculated from the simulated outcomes discussed above.

The values of these diameters were calculated from the simulated outcomes discussed above. ImageJ was used to measure the plastic deformation obtained at the end of the simulations. The measured values of these diameters at the end of 1st, 5th and 8th cycle of indentation are shown in Figure 14 and Figure 15.0.70.60.5r e t e m0.4i l l iM0.30.20.100102030405060708090X10^3 elementsDt Cycle 1Dt Cycle5Dt Cycle 8Figure 14.

Variation in the measured value of dt with increasing mesh size.0.60.5s r e t e m0.40.3i l l iM0.20.100102030405060708090X10^3 elementsDp Cycle 1Dp Cycle5Dp Cycle 8Figure 15.Variation in the measured value of dp with increasing mesh size.©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.1362An overlay of experimental and simulated load vs. depth of indentation curve is shown in Figure 16. A maximum indentation depth of 24% of the indenter radius and an unloading extent of 20% of the peak load at the end of each cycle were used during the simulation and experiment. The spine of the simulated load vs. depth of indentation curve matched very well the experimental data. However, the unloading sections of the curves were not a good match.

The spine of the simulated load vs. depth of indentation curve matched very well the experimental data. However, the unloading sections of the curves were not a good match. This mismatch will yield different estimates of chordal diameter and in extension stress, strain, yield, and ultimate strengths. The reason for the mismatch is suspected to be the lack of back stresses in the simulation.500400)N( d a o L300200100000.010.020.030.040.050.060.070.080.09Depth of indentationSimulatedExperimentalFigure 16. Overlay of experimental and simulation load vs. depth of indentation plots.

Overlay of experimental and simulation load vs. depth of indentation plots. A 20k node mesh with a peak indentation depth of 24% of indenter radius and a 20% unload at the end of each cycle was compared.ConclusionsThe ABI analysis consistently under-predicted the yield strength values through indentationThe Holloman’s power law was observed to be an invalid fit for true stress-true strain curve at low values of strainThe Voce formulation was observed to be a good fit for estimating the entire flow curve. • The ABI estimates for ultimate strength were close to the experimentally measured values • The load vs. depth of indentation simulated using finite element methods, overpredicted the plastic deformation during the indentationAcknowledgmentsThis work has been funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy.

DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or©2018 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.13630.1reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes. The supports from National Science Foundation Grants CMMI-1625736, and the Intelligent Systems Center (ISC) at Missouri S&T are greatly appreciated.Furthermore, the authors would like to acknowledge Kyle Stagner, Ian Wille, and BrianBullock for their aid in this research.References[1]S.-W. Khoo, S. Karuppanan, C.-S.

crease of Heat Transfer to Reduce Build Time in Rapid Freeze PrototypingMing C. Leu 1, Sriram P. Isanaka 1, Von L. Richards 2 1Dept. of Mechanical and Aerospace Engineering, Missouri University of Science and Technology, Rolla, MO 65409, USA 2Dept. of Materials Science Engineering, Missouri University of Science and Technology, Rolla, MO 65409- 0330, USA {mleu@mst.edu, sihyd@mst.edu, vonlr@mst.edu}ABSTRACTReduction of part build time in the Rapid Freeze Prototyping (RFP) process, which fabricates a 3D ice part layer-by-layer by depositing and freezing water droplets, has been achieved by increase of heat transfer. Three mechanisms have been experimentally investigated: 1) cooling the substrate, 2) use of forced convection, and 3) use of a chilling plate. Cooling the substrate is effective for parts of small heights but becomes ineffective with increase in part height.

Cooling the substrate is effective for parts of small heights but becomes ineffective with increase in part height. Forced convection produced desirable reduction in part build time but with the undesirable formation of frost on the built ice part. The use of chilling plate to increase heat conduction proved to be most effective. To ensure that the frozen ice from the deposited water can be easily removed from the chilling plate, various surface coats were investigated and the most effective surface coat was found to be a thin Teflon film. After incorporating the chilling plate we have successfully achieved 75% reduction in part build time.1. INTRODUCTIONThe Rapid Freeze Prototyping (RFP) process is a freeform fabrication process which builds a part by depositing water droplets and freezing them rapidly layer by layer [1-4]. The RFP apparatus, shown in Figure 1, consists of a substrate table capable of X-Y travel and a nozzle head capable of motion in the Z direction.

The RFP apparatus, shown in Figure 1, consists of a substrate table capable of X-Y travel and a nozzle head capable of motion in the Z direction. The entire apparatus, including the moving axes, is housed in a freezer that provides a temperature controlled environment. The required part geometry is traced by controlling the movement of an X- Y table while fine water droplets are deposited through a nozzle in a workhead mounted on the Z axis. The water droplets freeze on contact with the substrate or the previously deposited layer, forming a 3D geometry according to a CAD model. Two nozzles are used, one for the deposition of water and another for support material. The support material (eutectic dextrose solution) is needed to help generate parts of complex geometry with features such as holes, overhangs, slant walls, etc. Many parts of various shapes have been built using the RFP process.

Many parts of various shapes have been built using the RFP process. Compared with other SFF processes, RFP is a low-cost and environmentally friendly process because it uses water as the build1material and inexpensive equipment to build parts. Other benefits of this process include fine surface finish, low energy use, and ease of pattern removal for casting and molding applications.Z ElevatorPumpFreezerMotor DriverPipeSubstrateNozzleXY TableFigure 1. The Rapid Freeze Prototyping (RFP) apparatusOne potential application of the RFP process is using the fabricated ice parts as patterns for investment casting [5,6]. Wax patterns have traditionally been used in investment casting to make molds [7]; however, they have several disadvantages, the most significant of which is shell cracking during de-waxing. The cracking leads to difficulties in the manufacture of castings with thin cross-sections.

The cracking leads to difficulties in the manufacture of castings with thin cross-sections. Our overall research objective seeks to improve the capability of the RFP process so that it can be used to make thin-walled ice structures. Such improvement could facilitate new development of the shell mold process and enhance the means of producing small, intricate precision components.This paper describes an attempt to lower the temperature near the part build surface during the water droplet deposition in the RFP process in order to increase heat transfer, reduce build time, and enable fabricating thin-walled ice structures. The increase in heat transfer was achieved by: (1) Conduction – This was done through two mechanisms: (a) lowering the substrate temperature with coolants and (b) using a chilling plate; (2) Convection – This was done by using a fan to create forced convection in the system.

These various mechanism and their results on heat transfer are detailed in the present paper.2. COOLING SUBSTRATEThe first mechanism we investigated to increase heat transfer was modifying the substrate to reduce its temperature by incorporating coolant passages. This modification, compounded with the use of liquid nitrogen, permitted cooling of the substrate to the extremely low temperature of -140oC. The reduction in the substrate temperature resulted in a significant reduction in the water solidification time, thus permitting the production of a thinner part in a shorter build time.The rate of heat conduction is proportional to the difference between the temperatures of the two media involved in the heat conduction [8]. Liquid nitrogen, which boils at -196oC, acts as a perfect heat sink when it comes in contact with the aluminum substrate at the freezer temperature of -20oC, permitting rapid heat transfer to occur from the substrate to the liquid nitrogen.

The substrate base plate modification increased the surface area of contact between the liquid nitrogen and the aluminum plate. The pocket on the side of the substrate base plate acts as a reservoir for the manual input of liquid nitrogen. The liquid nitrogen flows into coolant passages drilled into the plate through the reservoir and cools the substrate from the inside. The heat transferred to the liquid nitrogen by the base plate changes the phase of the nitrogen from liquid to its more stable gaseous form, which vents through the miniature coolant passages built into the top of the base plate, thus enabling very low temperatures and very high rates of cooling. Figure 2(a) & (b) show the original substrate and the modified substrate, respectively.To measure the temperature achieved by the two substrates, nine holes were drilled into the substrates just under the substrate surface.

Through these holes, thermocouples were inserted at nine different points on each of the two substrate surfaces. Figures 3 and 4 plot the measured temperatures for the original and modified substrates, respectively. Because the temperatures measured by the nine thermocouples are very similar, only the temperatures obtained from two thermocouples are shown in these figures. With the original substrate, the thermocouples indicate a downward temperature trend (Figure 3), which is expected because of the use of liquid nitrogen. Even with considerable liquid nitrogen the lowest temperature achieved was only -40oC, after which the temperature leveled out and remained constant. However, with less than half the liquid nitrogen used in the original substrate, the temperature in the modified substrate dropped to -140oC before leveling out.(a)(b)Figure 2.

However, with less than half the liquid nitrogen used in the original substrate, the temperature in the modified substrate dropped to -140oC before leveling out.(a)(b)Figure 2. The original substrate (left) and the modified substrate (right))C(e r u t a r e p m e T)C( e r u t a r e p m e T02040Temperature data from thermocouple 160Temperature data from thermocouple 28010016011201180124013001Time (s)Figure 3. Temperature during part build (with the original substrate)2040Temperature data from thermocouple 160Temperature data from thermocouple 2801001201400100020003000 Time (s)40005000Figure 4. Temperature during part build (with the modified substrate)36016000Simulations performed in Fluent [9] to predict the time required to freeze a layer of water showed a substantial reduction in the water droplet freezing time with decrease in the substrate temperature.

The total heat loss, the liquid fraction, and the temperature decrease in the deposited water as functions of time were also simulated in Fluent to identify the water freezing time and the wait time between layers. The results obtained in Table 1 are for the first layer of water deposited on the substrate at various substrate temperatures. These results indicate that with an ambient temperature of -140oC, the wait time between layers can be shortened to about 5 seconds.These simulation results were used to guide experiments in which test parts were built with reduced wait time between layers. From Table 1 we conclude that the lower the substrate temperature, the shorter the time to freeze and the shorter the wait time between layers. The wait time was taken as the duration for the difference between the temperature of the deposited water and the temperature of the ambient to reduce to 1% of the difference between the initial water droplet temperature and the ambient temperature.

With the new substrate design, our experimental results showed that we could reduce the wait time between layers to about 15 seconds. With the original substrate, the shortest possible wait time was 40 seconds. The new substrate therefore permitted a reduction of over 60% in wait time between layers.Table 1. Simulation results (from Fluent) of time to freeze the first layer of water above the substrate and wait time at various ambient temperaturesTemperature (oC) -140 -120 -100 -80 -60 -40 -20 -10Time to Freeze (s) 0.8 0.9 1.1 1.0 1.5 2.5 4.7 9.6Wait time (s) 5.0 5.5 6.4 7.0 8.0 10.5 13.7 20.23. FORCED CONVECTIONWhen a layer of water is deposited on the aluminum substrate it is subjected to heat conduction through the substrate on one side and heat convection on its top surface. The high thermal conductivity of aluminum leads to effective heat transfer from the water layer at the very beginning of the ice part fabrication process.

The high thermal conductivity of aluminum leads to effective heat transfer from the water layer at the very beginning of the ice part fabrication process. As the part height increases, heat dissipation slows rapidly due to the low thermal conductivity of ice. A series of simulations were run in Fluent; the simulation results on the time taken for a deposited water layer to freeze at the ambient temperature of -35oC are summarized in Table 2. Clearly a layer of water deposited on an aluminum substrate transfers its heat much faster than a layer of water deposited on ice because ice has a much lower heat transfer coefficient than aluminum.As the height of the wall increases, there is greater tendency for the newly deposited water layer to transfer its heat to the ambient air through convection. Therefore, forced convection through fan blowing was tested as a means to increase heat transfer.

Therefore, forced convection through fan blowing was tested as a means to increase heat transfer. Areal-time data acquisition system was used to measure the temperature in the RFP apparatus using two thermocouples, one placed at 10 mm from the part build surface (thermocouple 1) and the other at 1 meter from the part fabrication region (thermocouple 2), as shown in Figure 5.Table 2. Simulation result of time to freeze for a water layer deposited at various heightsHeight 0 mm (on substrate surface) 1 mm 2.5 mm 5 mm 10 mm 12.5 mm 15 mm 17.5 mmTime to Freeze (s) 2.8 5.8 8.4 10.1 12 14.4 17.1 20.6Thermocouple 1Thermocouple 2Nozzle HeadSubstrate and tableFigure 5. Thermocouples placementThe measured temperatures are shown in Figures 6 and 7. From Figure 6 we can see that the heat transfer is concentrated in the neighborhood of the deposited water. This is evidenced from the fact that thermocouple 2 exhibited nearly constant temperature throughout the build cycle.

This is evidenced from the fact that thermocouple 2 exhibited nearly constant temperature throughout the build cycle. Liquid nitrogen was used to reduce the temperature around the part. The effect of pouring liquid nitrogen is indicated by the sharp downward spikes in Figures 6 and 7.When a similar investigation was conducted with the introduction of forced convection by using a fan to blow air cooled by liquid nitrogen into the build area, the cooling of the part was found to be much faster than without forced convection. PartsFreezerwere built successfully with a wait time of 20 seconds in this case, whereas without forced convection a wait time of 40 seconds was necessary. This reduction in wait time led to a 50% reduction in part build time. Although this cooling method provided considerable advantages, however, it also presented the problem of frosting. Frost appeared on the surface of the ice part as shown in Figure 8 when forced convection was used.

Frost appeared on the surface of the ice part as shown in Figure 8 when forced convection was used. Therefore, despite the reduction in wait time, this method did not appear to be a viable option for reducing the build time of ice parts.20)CThermocouple 2(e r u t a r e p m e T2530Thermocouple 135150011000115001Time (s)Figure 6. Temperature vs. time for part building without forced convection10Thermocouple 2)C15(e r u t a r e p m e T2025Thermocouple 1303501000020000Time (s)Figure 8. Temperature vs. time for part building with forced convection(a)(b)Figure 8. An ice wall built (a) without forced convection and (b) with forced convection4. CHILLING PLATEThe use of a pre-cooled plate to more effectively freeze deposited water was next investigated. The idea is to lower a chilling plate mounted on the z-axis after the deposition of each water layer to remove heart from the water and rapidly freeze it into ice.

The idea is to lower a chilling plate mounted on the z-axis after the deposition of each water layer to remove heart from the water and rapidly freeze it into ice. Figure 9 shows a CAD model and the actual chilling plate used. Besides cooling, the chilling plate provided a flat surface that could be used to flatten the newly deposited layer of water. Since conduction is a more effective mode of heat transfer than convection, it was conjectured that this plate could reduce the build time significantly. Preliminary experiments attempting to freeze a layer of water between the ice substrate and the chilling plate, however, indicated a tendency for the water to freeze onto the chilling plate. The ice part broke during its separation from the chilling plate. Therefore, identification of an interface film that would prevent the water from freezing onto the chilling plate became critical to the success of this technique.(a)(b)Figure 9.

Therefore, identification of an interface film that would prevent the water from freezing onto the chilling plate became critical to the success of this technique.(a)(b)Figure 9. The chilling plate investigated: (a) the CAD model and (b) the actual plateFluent was used to simulate the temperature and the liquid fraction in the chilling plate concept. The preliminary design of the chilling plate was a 3”×3” square with 0.2”in thickness. The temperature of the chilling plate increased from -20oC (253 K) to -8oC (265 K) after 6 seconds. The result is shown in Figure 10. Fluent simulation predicted that continuous use of the chilling plate for a period of over 50 seconds would result in residual heat buildup which will render the chilling plate ineffective. In an effort to prevent this heat buildup, instead of using a solid plate, the chilling plate was fabricated with coolant passages similar to the substrate plate.

In an effort to prevent this heat buildup, instead of using a solid plate, the chilling plate was fabricated with coolant passages similar to the substrate plate. Figure 11 shows the chilling plate along with the aluminum substrate plate.Residual heat buildupChilling plateLayer of waterSubstrateFigure 10. Temperature at time of 6 seconds after an aluminum chilling plate is in contact with deposited waterChilling PlateAluminum SubstrateFigure 11. The chilling plate and aluminum substrate50o44o(a)(b)70o65o(c)(d)80o105o(e)(f)Figure 12. Enhanced images of a water droplet on various surfaces: (a) no surface coat (polished face), (b) paint coated aluminum sheet, (c) mylar, (d) cellophane tape, (e) latex and (f) TeflonA major problem with the chilling plate approach was finding a surface coat that could prevent water from adhering to the chilling plate. The water repulsive properties of several materials were tested.

The water repulsive properties of several materials were tested. These materials included paint coated aluminum sheet, mylar, cellophane tape, latex, and Teflon. The best material was chosen by examining the contact angle between a water droplet and the surface of each material. The larger the contact angle, the greater is water’s tendency to repel that surface. Contact angles weremeasured using high-resolution digital photography. With ImageJ software [10], the contact angle between a water droplet and each surface coat was measured. The results are shown in Figure 12 and given in Table 3. It is clear that the largest contact angle occurs when the water droplet is deposited on the Teflon surface, of which the contact angle measured was 105o. This value of the contact angle is confirmed by measurements done by other researchers [11]. Therefore, Teflon was identified as the best surface coat.Table 3.

This value of the contact angle is confirmed by measurements done by other researchers [11]. Therefore, Teflon was identified as the best surface coat.Table 3. Contact angles of water droplets with various coat materials.Coat material No coat (polished face) Paint coated aluminum sheet Mylar (Transparencies) Cellophane tape Latex TeflonContact angle of water droplet (degrees) 50 44 70 65 80 105The chilling plate coated with a Teflon film of thickness 0.1 mm was then used to produce ice parts with a reduced wait time between layers. With the chilling plate, parts were built successfully with wait time of 10 seconds between layers. This is consistent with the wait time predicted by Fluent and it translates to 75 percent reduction in part build time. The significant reduction in wait time is a great advantage of this method of heat transfer, which can produce better ice parts than using forced convection because of no frost formation.5.

The significant reduction in wait time is a great advantage of this method of heat transfer, which can produce better ice parts than using forced convection because of no frost formation.5. CONCLUSIONThe research result presented has shown that by increase of heat transfer, much shorter wait times between build layers are possible than what could be achieved previously in the Rapid Freeze Prototyping process. This greatly reduces ice part build time and also ensures that thin-walled ice structures could be built. Three techniques were investigated to increase heat transfer and reduce build time in RFP fabrication of ice parts. Modifications made to the substrate by the use of coolant passages filled with liquid nitrogen to decrease the substrate temperature were successful in bringing about a significant reduction in part build time for the initial layers of water deposited on the aluminum substrate.

However, the effect of lowering the substrate temperature on reduction of wait time between layers decreased with increase in part height. Forced convection implemented with the use of a fan to circulate cold air produced desirable reduction of build time but with undesirable formation of frost. Very significant reduction in part build time was achieved using a chilling plate to increase the rate of cooling of deposited water by conduction. Contact angle measurement was used to identify the ideal surface coat, which was found to be a thin Teflon film. Fluent simulations predicted that the use of a chilling plate coated with a 0.1 mm Teflon film would bring about 75% reduction in part build time, which was confirmed with experiments.6. ACKNOWLEDGMENTThe research was funded by a Small Business Technology Transfer (STTR) grantfrom the National Science Foundation (IIP-0637556).BIBLIOGRAPHY1. Zhang, W., Leu, M.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceINCREASING INTERLAMINAR STRENGTH IN LARGE SCALE ADDITIVE MANUFACTURINGAlex Roschli*, Chad Duty*, John Lindahl*, Brian K. Post*, Phillip C. Chesser*, Lonnie J. Love*, and Katherine T. Gaul*Oak Ridge National Laboratory, Manufacturing Demonstration Facility, 2370 Cherahala Blvd, Knoxville, TN 37932AbstractInterlaminar strength of extrusion-based additively manufactured parts is known to beweaker than the strength seen in the printed directions (X and Y). With Big Area Additive Manufacturing (BAAM), large parts lead to long layer times that are prone to splitting, sometimes referred to as delamination, between the layers. Fiber filled materials, such as carbon fiber reinforced ABS, are used to counteract the effects of thermal expansion by increasing the strength in the X and Y directions.

Fiber filled materials, such as carbon fiber reinforced ABS, are used to counteract the effects of thermal expansion by increasing the strength in the X and Y directions. These fibers stay in-plane meaning that no fibers span from layer to layer, which would help counteract the weak interlaminar strength that causes splitting. A solution to this is a patent pending approach called Z-Pinning. The process involves strategically positioning voids across multiple layers that are backfilled with hot extrudate. This paper will explore the benefits and results of using Z-Pinning in large scale additive manufacturing.Introduction3D printing with the fused deposition modeling (FDM) process, works by adding material layer by layer to build a part. When FDM is done with thermoplastics, hot material is added, or extruded, on top of a previous printed layer that has had time to cool. This difference in temperatures between layers causes a weakness in the bond between these layers.

This difference in temperatures between layers causes a weakness in the bond between these layers. As layers grow larger, layer times also grow longer. This increase in layer time means that layers have more time to cool before the next layer is extruded on top. This results in an increased temperature gradient across the part. As thermoplastics cool, they contract. This cooling and contracting creates residual stresses throughout the printed part. If the stresses become too much, delamination, or splitting of the layers, occurs. The world’s first 3D printed car, Strati, had this issue as seen in Figure 1. Strati had layer times exceeding an hour which allowed the previous layer to cool to near room temperature before being extruded over.This manuscript has been authored by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy.

DE-AC05-00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes. The Department of Energy will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).543Figure 1: An early test print of the Strati that resulted in numerous delaminations that had to be glued togetherTo attempt to fix the problem of the delamination for large prints, Oak Ridge NationalLaboratory (ORNL) started looking for methods of increasing the strength in the Z-direction and improving the bond layer to layer.

A process called Z-Pinning was developed that works by leaving a hole, or holes, in each layer and then coming back along and extruding material into the hole to fill it [1]. This allows hot material to span several layers at one time thus allowing the polymer chains to go between layers and not just stay within the layer plane (Figure 2). When using fiber reinforced materials, the fibers now span the layers which gives added strength in the Z-direction. This paper will explore the preliminary testing of Z-Pinning with BAAM.Figure 2: Cross-section showing pins in green544Z-Pinning Tests on BAAMTesting on BAAM started with simple, single pin prints. This was done to determine the best pin size. A pin that is too small won’t properly fill the void, whereas a pin that is too large will take too long to fill. Another issue is a pin that is too tall; a pin that is too tall won’t fill to the bottom. To test these variables, holes of various sizes were printed and tested at one time.

Another issue is a pin that is too tall; a pin that is too tall won’t fill to the bottom. To test these variables, holes of various sizes were printed and tested at one time. The holes were cut in half so that a piece of clear polycarbonate could be placed in front and video could be captured. The testing showed that a hole of diameter four times that of the nozzle diameter filled best and that eight layers was the best height.Figure 3: A print done to find the optimal hole size. Outer holes show the polymer didn’t reach the base of the hole.After the appropriate hole size was found, full pieces were made with one single centralpin. This was done to ensure that the pin filled appropriately and didn’t have negative effects, like overfilling or swelling, on the surrounding area. These parts were then cross-sectioned to ensure a fully dense pin was created (Figure 4).Figure 4: Two cross-sectioned parts with one central pin.

These parts were then cross-sectioned to ensure a fully dense pin was created (Figure 4).Figure 4: Two cross-sectioned parts with one central pin. The pin is not distinguishable because it is fully dense.Multi-pin test pieces were printed after the single pin pieces. These were done using an“8-4” pinning scheme. This means eight-layer tall pins that start every four layers, which is also545known as an A-B approach (Figure 5). These parts were five inches by five inches and thirty inches long. Five sets of four test pieces were made with each of the sets representing a different fill pattern. To test the sets against each other, they were loaded in a three-point bending machine, which will be discussed in the next section.Figure 5: Overhead view of the AB Pinning approachThe first set was a control set where the infill pattern was fully dense (Figure 6). Theparts were printed with one outer contour and a horizontal raster infill that rotated by 90 degrees each layer.

Theparts were printed with one outer contour and a horizontal raster infill that rotated by 90 degrees each layer. The second set was made in a similar fashion but with a sparse infill pattern (Figure 7). The infill spacing for this was set at 0.45in, which results in a density of 75%. The resulting part would have a similar weight to the third set, which had five pins. Having a similar weight would allow for a direct comparison of strength to weight.Figure 6: Three-Point Bend set 1, fully dense infill.546Figure 7: Three-Point Bend set 2, sparse infill with 75% density.As previously mentioned, set three had five pins in an 8-4, AB, pinning style. The pinsalong one diagonal were the A pins and the remaining two pins, which were on opposite corners, were the B pins.

The pinsalong one diagonal were the A pins and the remaining two pins, which were on opposite corners, were the B pins. A G-Code rendering of a single layer where three pins are filled is shown in Figure 8.Figure 8: Three-Point Bend set 3, five pins in a 4-8 style.547Figure 9: Short test version of set 3The fourth set was made without pins. It was made using a special infill pattern designed to mimic the geometry of set three. This was done by creating tiny holes at the centers of the pin locations and forcing the slicer to print around the holes, Figure 10.Figure 10: Three-Point Bed set 4, infill that mimics the geometry of set 3The remaining set of test beams were created using a custom infill pattern with four pins. The pins in set 3 were circular with a perimeter bead surrounding them (Figure 8). This caused a non-uniform infill pattern surrounding the pins that ultimately made the layer take more time and caused a lot of voids.

This caused a non-uniform infill pattern surrounding the pins that ultimately made the layer take more time and caused a lot of voids. The combination of these two factors made the layer to layer bond very weak. A suitable solution was found that four pins could be placed at the corners in square holes with the remainder of the volume being filled with a raster infill pattern (Figure 11). This meant a better layer to layer bond and less starts and stops causing a shorter layer time.548Figure 11: Custom infill pattern with four pin locations in the cornersThree-Point Bend TestingThe initial small-scale research for Z-Pinning involved material strength testing using tensile bars and a pulling machine [2]. For large-scale, it was challenging to determine a print that could be properly used to measure tensile strength. Tensile bars need to be cut in a way that pins are fully included without being cut with the tensile bar spanning across all layers.

Tensile bars need to be cut in a way that pins are fully included without being cut with the tensile bar spanning across all layers. This would have made for very large samples to pull. An alternate solution to tensile samples was a three-point bend test. A three-point bend test involves placing a beam that is supported on the ends and loaded in the middle [3]. Large beams could be printed with many pins inside, then broken on a large bending machine to measure the load.The printed beams were 5in x 5in x 30in each. The printed height was 30in so that pinswould be stacked on top of each other. This also allowed the samples to be laid flat in the testing rig so that the load would be applied between the layers and force the pins to break. Figure 12 shows an example of the testing setup with one beam loaded in the bending machine.549Figure 12: Three-point bending machine setupResultsFor each sample set, four bars were broken on the three-point bending machine.

Fromthat data, the average was taken for each set and can be seen in table 1. The highest bending load was seen in set 4. This is because the infill pattern created an increased layer to layer bond strength by maximizing the contact area from layer to layer (Figure 13), even in comparison to solid infill (Figure 14). This was also the result of the part being overfilled causing more material to be packed into each layer.Set 1 Set 2 Set 3 Set 4 Set 5Load (lbs) 5089 3698 4252 8984 5585Table 1: Max loading weight for each beam sample set550Figure 13: Set 4 showing the increased contact area between layers.Figure 14: Set 1 with less contact area than set 4.551Solid and sparse line infill patterns, like sets 1 and 2, are most commonly printed on largescale machines. When comparing the two sets to the improved pin set, set 5, a strength increase is seen for the pinned set.

When comparing the two sets to the improved pin set, set 5, a strength increase is seen for the pinned set. The pinned set increases the strength by 9.75% over the solid infill and 51.03% over the sparse infill (Table 2).Table 2: Comparison of solid (set 1), sparse (set 2), and pinned (set 5) maximum bending loads.Non-Cylindrical PinsAs seen in the results of the three-point bend testing, pins don’t always shear during testing and sometimes break off at the boundary between pins (Figure 15). This causes the pin to be pulled out of the part and remain intact. The force to break a pin at the layer boundary is less than the force required to shear the pin in half causing a lower load strength for the beam.552Figure 15: Fracture of a three-point bend sample. The bottom right pin illustrates a pull-out.To fix the issue of pins pulling out, testing was done with pins that are non-cylindrical, or have flanges (Figure 16).

The bottom right pin illustrates a pull-out.To fix the issue of pins pulling out, testing was done with pins that are non-cylindrical, or have flanges (Figure 16). These non-cylindrical pins expand at the bottom such that they cannot be pulled up out of the part. If the cylindrical section has a diameter equal to that of the nozzle, fibers should be aligned in the Z-direction with the extrudate. Testing of samples with this pin geometry has not happened yet but is upcoming.Figure 16: Top, CAD model of a test piece with non-cylindrical pins. Bottom, printed sample pieces without pins filled.553Future WorkThis paper and the research it entailed is just the beginning of the testing and development of Z-Pinning for BAAM and other large scale additive manufacturing systems.

Future work will include Z-Pinning during infill, automating the pinning process with slicing software, mixing of materials, and testing hot pins.In this paper, all the pins were specifically designed into the CAD model. The slicerwould leave the pin area as a void that could be later filled with a post-processing script. Future development will remove the need to design the pins in CAD. Instead the pins will be placed within the infill sections using a grid-based infill (Figure 17). The goal is to fully automate this within ORNL Slicer.Figure 17: Example of a Z-Pin within a section of grid infillMixing materials will allow the main structure of the part to be made with one material,while the pin is made with a separate material. This will provide the opportunity to use a pin material that has a higher fiber loading, which would increase strength in the Z-direction. It will also allow for dissimilar materials to be used for the pins to help the layers bond together better.

It will also allow for dissimilar materials to be used for the pins to help the layers bond together better. Low melt strength polymers, such as HDPE, can be used to help tightly fill the pin voids. Mixing materials leads to the idea of “hot pins,” which are pins extruded at a higher temperature to help better fill the void. Hot pins can be done using the same material or a dissimilar material.ConclusionZ-Pinning is a new approach to solving the issue of poor interlaminar strength with FDM parts. Small scale testing has proven that Z-Pinning is effective for increasing the bond between layers, and this paper documented very early research into its application with large scale 3D printing. With Carbon Fiber ABS, an increase in layer to layer strength was seen by applying the Z-Pinning technique. A 9.75% increase was achieved in comparison to a part with solid infill and a 51% increase was found in comparison to a part with 75% dense infill.

A 9.75% increase was achieved in comparison to a part with solid infill and a 51% increase was found in comparison to a part with 75% dense infill. This is just the beginning of testing and applying Z-Pinning to BAAM and other large scale processes, but the results show promise for printing parts with improved Z-strength.554AcknowledgementsThis material is based upon work supported by the U.S. Department of Energy, Office ofEnergy Efficiency and Renewable Energy, Office of Advanced Manufacturing, under contract number DE-AC05-00OR22725.References[1] C. Duty, J. Failla, S. Kim, J. Lindahl, B. Post, L. Love, and V. Kunc. “Reducing Mechanical Anisotropy in Extrusion-Based Printed Parts.” Proceedings of the Solid Freeform Fabrication Symposium, Austin, TX (August 7-9, 2017).[2] Vertical Pins for Strengthening Extrusion-Based Printed Parts.” SAMPE 2018, Long Beach, CA (May 21-24, 2018).S. Kim, T. Smith, J. Failla, J. Lindahl, V. Kunc, and C. Duty.

Kim, T. Smith, J. Failla, J. Lindahl, V. Kunc, and C. Duty. “Parametric Analysis on[3] Reinforced Plastics and Electrical Insulating Materials, ASTM International, West Conshohocken, PA, 2017, www.astm.orgASTM D790-17, Standard Test Methods for Flexural Properties of Unreinforced and555.

Moreover, for additively fabricated parts consisting of Ti6Al4V, hot isostatic pressing(HIP), is widely used as a post treatment in order to increase the relative density of the built parts.However, one aspect which has rarely been addressed so far, is the increase of process speed,accompanied by a reduced relative density in the as-built condition and a subsequent hot isostaticpressing process to achieve the desired dense material. The approach here is to use the standardprocess route as described, but intentionally increase the process speed and accept a certain valueof porosity.

The approach here is to use the standardprocess route as described, but intentionally increase the process speed and accept a certain valueof porosity. The focal objective of this study is the identification of a parameter-set with the highestpotential for an increase of process speed and subsequently reduce the internal defects during thehot isostatic pressing process to achieve completely dense components.IntroductionAdditive manufacturing (AM) processes are widely used in industry and represent a regularfeature of the product development process [1]. There are different technologies (i.a., selectivelaser melting of metals, selective laser sintering of polymers and fused deposition modelling ofpolymers), which are summarized under the term AM. All of them create a part by adding materialrather than subtracting material as conventional machining [2, 3].

All of them create a part by adding materialrather than subtracting material as conventional machining [2, 3]. The huge possibilities of thistechnology, in case of design freedom for structural and functional parts, lead to a high demand forlightweight components in several industries (i.a. aerospace, medicine, automotive) [2].Concerning SLM, it is a tool-free process in which the component is generated in a layer by layer1506fashion. Therefore, the components three dimensional geometry design is digitally sliced in severaltwo dimensional layers, which are loaded in the machine. Thin powder layers, e.g., 50µm, aredeposited on a building plate one after another with a recoater and the two dimensional geometryis successively molten by the ytterbium laser. After solidification of the current layer, the buildingplate is lowered in z-direction, in exactly one layer height, and the recoater deposits the next powderlayer such that the process restarts.

After solidification of the current layer, the buildingplate is lowered in z-direction, in exactly one layer height, and the recoater deposits the next powderlayer such that the process restarts. Thereby nearly any complex geometry can be manufacturedstep by step [3, 4]. The unmolten material can be reused after a sieving process. Due to the highenergy input, large thermal gradients and emerging residual stresses occur during solidification.Thus, each component requires support structures to avoid warping. [4, 5].Nowadays, the highest cost driver is the SLM-processing time and the mechanical post-processing.On that account, SLM machine manufacturers develop enhancements to increase the efficiency oftheir machines. SLM machines are available with multi-laser optical systems to increase processspeed and create fully dense components [6, 7].

SLM machines are available with multi-laser optical systems to increase processspeed and create fully dense components [6, 7]. Current machines as the single laser or the multi-laser machines can achieve building rates from 20cm³/h up to 105cm³/h, by employing a laserpower between 200W and 1kW [7]. With the help of multi-laser systems the building rate can beincreased by the amount of lasers as a factor in building time reduction of a similar build job. Thesesystems divide the 2D building area in separate scanning areas [3].

Thesesystems divide the 2D building area in separate scanning areas [3]. Another field of researchaddresses high power SLM, which means to increase the laser power up to, e.g., 2kW and adjustthe scanning strategy in order to generate components with a density of more than 99.5% [8, 9].An established processing route for SLM components consisting of Ti6Al4V can be divided in fivesteps: the production process; a first heat treatment (stress-relief annealing, in which componentsstay on building plate); mechanical post-processing (cut components from building plate, andremove support structures); a second heat treatment (hot isostatic pressing to reduce internaldefects); and a final mechanical post-processing of interfaces and functional surfaces (i.a., millingor further machining) [3].

During hot isostatic pressing, which is a process where components arepretermitted high temperatures up to 1200°C and isostatic pressure around 1000bar, internal defectsare closed and density increase up to 99,99%. The conditions during the HIP process lead tolocalized plastic deformation and solid-state diffusions take place [6].Within the current work, the focus is on the powder-bed based selective laser melting (SLM) ofmetals, respectively titanium alloy Ti6Al4V. Especially, in the aerospace industry, Ti6Al4V1507components fabricated by SLM, offer high potential due to its severe weight savings potentialbased on the low density 4,43 g/cm³ of Ti6Al4V.Therefore, the approach here is to decouple the increase of process speed from the hardware of amachine, but develop an adapted process idea.

The aim is to use the default and known processroute for Ti6Al4V parts, identify a high speed SLM parameter set and get the same density asbefore through the use of the HIP process.The approach is to optimize the first step in the described process route by optimizing the buildingparameters and accept a slightly higher porosity due to an increased process speed. During the HIPprocess internal defects are eliminated. Regardless of whether the specimens have a density valuearound 99,9% or around 97%. The results show, that the density at the end of the process route canbe similar, the only difference is the building time of the specimen. The building time can bereduced up to 50%, while still getting dense products.

The building time can bereduced up to 50%, while still getting dense products. Due to the use of a default and known processroute with little changes in the subject, the procedure described in here is transferrable to severalother systems as well.Experimental DetailsThe material used in this work was gas atomized titanium alloy Ti6Al4V powder. Thechemical composition of the powder was determined by different testing methods, as can be seenin Table 1.Table 1 – Chemical composition measured of the used Ti6Al4V powder.Chemistry AlCFeHNOTiVaOtherTestICPLecoICP-LecoLecoLecoDiffICPICP-MethodMSMSResult6.360.010.040.00440.020.12Balance 3.980.16The size of the atomized particles were in the range from d(0.1): 19.8µm – d(0.9): 77.15µm withan average particle size of d(0.5): 41.99µm. As shown in Figure 1 a), the particle size distributionpossessed a Gaussian function.

As shown in Figure 1 a), the particle size distributionpossessed a Gaussian function. In addition, the particle morphology can be described as sphericalin shape (Figure 1 b)). The particle morphology was analyzed with a scanning electron microscope1508(SEM). A spherical particle morphology was required for a sufficient powder flowability and thusa homogenous powder layer deposition.Figure 1 – Particle size distribution and particle morphologyFor specimen fabrication, a SLM 280HL (SLM Solutions Group AG, Lübeck, Germany) wasapplied. The system was supplied with a 400W Ytterbium fiber Laser. The beam focus diameterpossessed a size of 70µm and a Gaussian power distribution. Table 2 reveals the reference SLMparameter-set, which was basis for the parameter variation. Cylindrical specimens with a heightand a diameter of 10mm were additively manufactured.

Table 2 reveals the reference SLMparameter-set, which was basis for the parameter variation. Cylindrical specimens with a heightand a diameter of 10mm were additively manufactured. Overall, the design of experiment included20 different parameter-sets and 60 specimens in total by the variation of laser power (275W up to385W), scanning speed (760 mm/s up to 2500 mm/s) and hatch distance (0,12 mm up to ,018 mm).The resulting volume energy density varied from 26,7 J/mm³ to 60,31 J/mm³. The parameter studybased on a previous investigation of the authors. All specimens were built with a layer thicknessof 50µm and a building-platform temperature at 200°C.

The parameter studybased on a previous investigation of the authors. All specimens were built with a layer thicknessof 50µm and a building-platform temperature at 200°C. In order to reduce internal defects, a HIPprocess was performed at 920°C/1000 bar on one half of the specimen.Table 2 – Reference parameter for specimen productionParameterLaserpowerScanning speedHatch distanceVolume energy[W][mm/s][mm]density [J/mm³]Reference2757600,1260,3Hatch1509For dark field optical images, the samples were mechanical grinded down to 6µm grit size and thenvibro-polished using a colloidal silica suspension with a pH-value of 10 and a particle size of0.06µm. The metallographic cross sections were subsequently analyzed by the software ClemexVision Lite (Longueuil, Canada) with regard to grayscale differentiation in order to quantify theprevailing relative density. Furthermore, the specimen were etched for 20 sec. in etchant accordingto Kroll.

Furthermore, the specimen were etched for 20 sec. in etchant accordingto Kroll. The building time calculation was done with the Software Magics from Materialise.Results and DiscussionFigure 2 depicts representative dark field micrographs of the samples’ cross sections (x-z-plane); three specimens in the as-built condition and similar specimens after the HIP process. Thedensity in as-built condition processed with the standard parameter-set is determined at 99,97%and at 99,99% after the HIP process. The samples fabricated with 29,17 J/mm³ and a calculatedprocessing time reduction of 51% possess a relative density in the as-built condition of 99,21% andof 99,95% after the HIP process.

The samples fabricated with 29,17 J/mm³ and a calculatedprocessing time reduction of 51% possess a relative density in the as-built condition of 99,21% andof 99,95% after the HIP process. The samples manufactured with 30,85 J/mm³ and a calculatedprocessing time reduction of 49% has a density in as-built condition of 99,14% and of 99,97% afterHIP process.As buildHIP processed60,31 J/mm³ - Reference60,31 J/mm³ - Reference151029,17 J/mm³29,17 J/mm³30,85 J/mm³30,85 J/mm³Figure 2 - Dark field optical images displaying cross sections of the SLM processed specimens in the as-built condition (a)-(c) as well as the HIP-condition (d)-(f).In Figure 3, the effect of the HIP process on the porosity and the resulting time reduction is shown.It can be observed that by employing a subsequent HIP process the density is significantly higherthan before the post-treatment. Sample 1 (S1) represents the reference parameter-set with a slightlyincrease of laser-power.

Sample 1 (S1) represents the reference parameter-set with a slightlyincrease of laser-power. The samples S5 and S8 are built with 29,17 J/mm³ and 30,85 J/mm³ andcan be compared to the cross section images shown above. The highest density increase wasmeasured for the condition S10, in which the porosity decreases from over 6% to less than 0.01%,while achieving a time reduction of approximately 52%. In addition, an economic approval was1511carried out. Based on three different buildjobs with different building heights, space utilization andvolume components, the potential was examined on economic aspects.Figure 3 – Optical density analysis of cross section images (x-z-plane)In Figure 4, optical images are displayed revealing six representative microstructures of the SLM-processed and HIPed specimens after etching. As Leuders et al. showed in previous investigations[10,11], a lamellar α-titanium phase can be detected in this images as well.

As Leuders et al. showed in previous investigations[10,11], a lamellar α-titanium phase can be detected in this images as well. Since the β-transustemperature of Ti6Al4V is at 995°C, there is no β-titanium expected.1512Figure 4 – Microstructure of Ti6Al4V specimens with different parameter setsIn consequence, the similarity of the microstructure leads to the assumption that the chosen volumeenergy density does not lead to any microstructural changes after HIP process. In combination withthe density measurements, it can be presumed that it is feasible to fabricate components with similarproperties in severely reduced SLM-processing time (time reduction between 13% and 48% inFigure 43) by conducting the investigated procedure.ConclusionsThe presented results demonstrate the potential of the high-speed SLM strategy.

A marginalprocess route modification can obviously lead to a severe reduction of 50% in SLM-processingtime.Based on the conducted investigations, the current findings can be summarized as follows:1. The HIP process is able to reduce porosity from up to 6% to 0.01%.2. Processing time can be reduced up to 50% on optimal conditions for Ti6Al4V componentsusing the described process route.3. By varying the parameter sets and the knowledge gained with regard to the density, it isnow possible to adjust the density in a targeted manner.15134. No negative modifications in the microstructure of the specimens after HIP process causedby the build time reduction and the accompanying build rate increase.The described approach has to be validated with experiments on the mechanical properties ofspecimen produced with the optimized parameters and the adapted process route.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceIncreasing the Interlayer Bond of Fused Filament Fabrication Samples with Solid Cross-Sections using Z-PinningChad Duty1,2, Tyler Smith1,2, Alexander Lambert1,2, Justin Condon1,2, John Lindahl2, Seokpum Kim1,2, Vlastimil Kunc1,2,31University of Tennessee 2Manufacturing Demonstration Facility, Oak Ridge National Laboratory1 3Purdue UniversityAbstractThe mechanical properties of parts made by fused filament fabrication is highly anisotropic, with the strength across layers (z-axis) typically measuring ~50% lower than the strength along the direction of the extruded material (x-axis). A z-pinning method has been developed in which material is extruded in the z-direction to fill intentionally aligned voids in the x-y print pattern.

A z-pinning method has been developed in which material is extruded in the z-direction to fill intentionally aligned voids in the x-y print pattern. In previous studies that involved a sparse rectilinear grid cross-section (35% infill), the z-pinning approach demonstrated more than a 3.5x increase in strength in the z-direction. The current study expanded these efforts to evaluate the use of z-pins in a printed sample with a solid cross-section. Although a solid cross-section is more common in structural components, it is much less forgiving of instabilities that may occur in the z-pinning approach (such as over-filling). Even though this study utilized a low pin volume (~43% fill factor), the pinning approach demonstrated a 40% increase in z-direction strength for solid samples that had similar printing times.IntroductionMaterial extrusion (MatEx) printing techniques deposit successive layers of material in the x- y plane to create three-dimensional objects.

For a simple 0/0 infill pattern, individual layers are comprised of long beads of continuous material (x-direction) that are a placed adjacent to one another in the y-direction. Various infill patterns are used to balance mechanical properties in the x-y plane, the most common of which are a 0/90 or a 45/-45 configuration in which the primary deposition direction rotates 90 degrees from the previously deposited layer. Regardless of the infill pattern, the layer-wise deposition results in relatively continuous material in the x-y plane, but discontinuous material in the z-direction (across layers). Transmitting a tensile load in the z- direction relies on the strength of intermittent bonds between overlapping sections of deposited beads in successive layers.This manuscript has been authored in part by UT-Battelle, LLC under Contract No. DE-AC05-00OR22725 with the U.S. Department of Energy.

DE-AC05-00OR22725 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a non-exclusive, paid-up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for United States Government purposes. The Department of Energy will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan).1769The discontinuous nature of MatEx structures leads to a significant level of mechanical anisotropy. A number of studies have shown that the tensile strength of parts produced on small scale printers with common materials, such as acrylonitrile butadiene styrene (ABS) and polylactic acid (PLA), is reduced by 50% to 80% in the z-direction [1-3].

The addition of reinforcing fibers to the polymer matrix increases the degree of mechanical anisotropy on both small- and large-scale platforms [1, 2, 4, 5].Various approaches have attempted to reduce the mechanical anisotropy of MatEx components. The majority of efforts improve the interlayer bond strength by controlling the temperature between layers to encourage molecular diffusion and increased entanglement. Localized heating of the bond interface during the printing process using forced air [6], laser irradiation [7, 8], and infrared lamps [9] have each demonstrated improved performance. Post- processing techniques have also attempted to improve bond strength of MatEx parts by exposing the component to ionizing radiation [3], microwave energy [10], or elevated temperatures [11].

Post- processing techniques have also attempted to improve bond strength of MatEx parts by exposing the component to ionizing radiation [3], microwave energy [10], or elevated temperatures [11]. Other researchers are exploring techniques to print out-of-plane to either align the strong direction of the structure with the primary load direction [12, 13] or print a secondary structure around the periphery of the part to reinforce the layered core [13, 14].The current research investigates the effectiveness of a z-pinning approach for MatEx printing, which deposits continuous material across layers throughout the bulk of a part by extruding into intentionally aligned voids. The patent-pending z-pinning approach that was introduced in 2017 [15, 16] is illustrated in Figure 1. As the primary structure (white) is being printed, voids (diameter, DH) are positioned such that they align in the z-direction.

As the primary structure (white) is being printed, voids (diameter, DH) are positioned such that they align in the z-direction. In the example of Figure 1, the first two layers of the primary structure are printed with voids (a) and (b) remaining. Between deposition of the second and third layers, the print head is positioned over void (a) and material is extruded while the print head is stationary, filling void (a). The second void (b) remains unfilled until layers three and four are printed along with the first two layers of void (c). After layer four is complete, the print head is held stationary over void (b) and material is extruded to fill the void14DH ~187(d)654(c)- -3l 20Seam - - - (a)(b)Figure 1. Schematic of the z-pinning process.1770Lp(note that the first two layers of void (c) remain open). After layers five and six are deposited, the pinning process is repeated to fill void (c), leaving a seam between pins (a) and (c).

After layers five and six are deposited, the pinning process is repeated to fill void (c), leaving a seam between pins (a) and (c). A similar seam exists between layers four and five at the interface between pins (b) and (d). Note that the seam positions are staggered due to the alternating pin positions.The pinning pattern can be extended throughout the bulk of the printed component in all directions. In the example of Figure 1, the standard pin length (LP) is four layers and pins are extruded into alternating holes every two layers. The labeling convention for this type of configuration is a “4-2 pin”. The theoretical volume of the void can be calculated from the void diameter (DH) and length (LP). The extruded volume of the corresponding pin to fill the void can range from a small percentage of this theoretical volume to well over 100%, depending on the surrounding structure.

The extruded volume of the corresponding pin to fill the void can range from a small percentage of this theoretical volume to well over 100%, depending on the surrounding structure. In an earlier study [17], sparse rectilinear structures were investigated that had an open architecture, allowing fill factors (pin volume / void volume) to reach 120%. The ultimate tensile strength in the x-direction (UTSX) and z-direction (UTSZ) are compared in Figure 2 for carbon fiber reinforced PLA (CF-PLA) structures evaluated during that study. The solid structure was a dense rectilinear grid pattern with no pins, which demonstrated an expected anisotropy of ~33%. The sparse rectilinear grid sample spaced the adjacent beads to create an open architecture with a 35% infill. The sparse structure without pins was much weaker than the solid structure (80-90% drop) and it demonstrated an even higher degree of mechanical anisotropy (~63%).

The sparse structure without pins was much weaker than the solid structure (80-90% drop) and it demonstrated an even higher degree of mechanical anisotropy (~63%). When 8-4 pins were added to the sparse structure, the UTSZ more than tripled and the mechanical anisotropy was cut in half.50DX-AxisII Z-Axis45rn c.. ~ .c 'li'o C ~ .... Vl ~ ·.;; C ~ .l!:l "' E ~ 1040353025201533% Drop30% Drop::::)50Solid Grid, No PinSparse, No PinSparse, 120% PinsFigure 2. Average ultimate tensile strength of CF-PLA printed samples.The current study extends the z-pinning technique to more conventional solid structures that have a 45/-45 infill pattern [18]. Placing z-pins into a solid structure is much more difficult due to the lack of interstitial space for the pin to expand into.

Placing z-pins into a solid structure is much more difficult due to the lack of interstitial space for the pin to expand into. Due to the layered nature of a MatEx structure, the internal walls of the voids are inherently rough (scalloped – see Figure 1), so the pin can still partially mechanically interlock with the surrounding structure. However, since the solid structure is less forgiving of print variations, it is difficult to achieve fill factors that approach 100%.1771Experimental MethodsThe material used for this study was a 15% carbon fiber reinforced PLA (CF-PLA) purchased as CarbonX from 3DXTech. Samples were produced on a MakerGear M2 desktop printer with a print temperature of 245ºC. The individual tensile samples were harvested from a single vertical wall, measuring 127 mm (x) x 12.7 mm (y) x 127 mm (z). The dimensions of the tensile sample are shown in Figure 3 as they are waterjet cut from the printed vertical wall.

The dimensions of the tensile sample are shown in Figure 3 as they are waterjet cut from the printed vertical wall. The samples were dried in an oven at 50ºC for at least 8 hours and stored in a desiccant chamber prior to tensile testing on an MTS servo-hydraulic test system (110 kN load cell, strain rate of 3 mm/min).R 76138.128.5L L...___.,I~ 31.8 ~t=c=--:2:_7_-1______ 140 - - - - - --Figure 3. Tensile sample dimensions (left) in mm waterjet cut from printed wall (right).A conventional solid sample with 45/-45 infill pattern was printed as an initial control reference. As shown in Figure 4, the print time for the solid sample included only sum of the times to print the individual layers (No Pin, No Pause). When pins were added on every fourth layer (8- 4 Pins), it consumed an additional 55 seconds each time, increasing the print time of that particular layer by more than 50%.

When pins were added on every fourth layer (8- 4 Pins), it consumed an additional 55 seconds each time, increasing the print time of that particular layer by more than 50%. Since the strength of the interlayer bond has been attributed to maintaining a high interfacial temperature [19-21], the additional time during every fourth layer could significantly cool the interface and weaken the structure. Therefore, a more appropriate control sample was also printed without pins that included a pause on every forth layer to match the print time of the pinned samples (No Pin w/ Pause). In practice, this could be representative of a much larger part that would naturally have longer layer times.1772No Pin, No PauseNo Pin w/ Pause8-4 Pins,e - - - - - - Pause Tim .,= 55 s,t - - - - - - - - -1 Layer Tim e ,___ ____ __, = 1 m 54 sFigure 4. Print time comparison among solid samples.Two pinning configurations were considered for this study.

Print time comparison among solid samples.Two pinning configurations were considered for this study. Both sample sets utilized three rows of 8-4 pins across the thickness of the sample with a fill factor of 43%. The pinning configuration within the x-y plane was varied among the two sample sets as shown in Figure 5. For this figure, consider that the “Kth” pin in the “A” location would be deposited after layer (Kn) while the Kth pin in the “B” locations would be deposited after layer (Kn+m), where pins of length (n) are inserted every other (m) layers. For the example of the 4-2 pins (n=4, m=2) in Figure 1, pin (b) would be the first “A” pin (K=1) deposited after layer 4, and pin (c) would be the first “B” pin (K=1) deposited after layer 6. When (n/2 = m), the seams between “A” pins are located on the same layer as the mid-point of the “B” pins. Figure 5 shows a “Standard” pinning configuration where the “A” pins are located on alternating rows of the x-y plane.

Figure 5 shows a “Standard” pinning configuration where the “A” pins are located on alternating rows of the x-y plane. If a fracture occurred across a given x-y plane, the Standard configuration involving only three rows of pins would be unevenly balanced. The “Staggered” configuration alternates the pin locations along a given row, providing more balance across a 3 row structure.StandardABABABABABA2/3 A'sAAAAAA1/3 B'sStaggeredABABAB1/2 A'sBBAABBAABBA1/2 B'sFigure 5. Comparison of pin spacing patterns in the x-y plane.1773Results and DiscussionThe ultimate tensile strength in the z-direction of the printed samples are shown Figure 6. The conventional solid sample that was printed without pauses is clearly the best performing structure, with a UTSZ of 33 MPa.

The conventional solid sample that was printed without pauses is clearly the best performing structure, with a UTSZ of 33 MPa. When the same structure was printed with a 55 second pause on every fourth layer, the strength dropped to a baseline value of 12 MPa (64% reduction). The addition of z-pins to the structure improved the UTSZ by roughly 40% regardless of the pinning configuration used (Standard = 17 MPa, Staggered = 16 MPa). Although the strength of the pinned structure did not approach the solid sample printed without pausing, this represents a significant improvement in strength when compared to parts that may be cooling down considerably between successive layers (e.g. a large part with a naturally long layer time).35.. - -DZ-Ax is30ro c.. ~ 25 ..r::. 1i'o C .... QJ ...... Vl ..!!! "' C ~ QJ ...... ro E ~ :::>201510I -64% I-· - -- - J 44% ~ - ...

- -DZ-Ax is30ro c.. ~ 25 ..r::. 1i'o C .... QJ ...... Vl ..!!! "' C ~ QJ ...... ro E ~ :::>201510I -64% I-· - -- - J 44% ~ - ... - - - - -T lT=-11~~~J l> - - ---- -> - - -5 ----0No Pin, No PauseNo Pin w/ Pause8-4 Pins, Standard8-4 Pins , StaggeredFigure 6. Average ultimate tensile strength of solid printed samples (CF-PLA).The fracture surfaces of the tensile samples are shown in Figure 7. It is apparent that the unpinned sample had a very clean fracture surface, with the crack front progressing unobstructed across the layer-to-layer interface (note: the two unpinned sample sets looked very similar). In contrast, the two pinned samples resulted in very tortuous fracture surfaces as the crack front deflected around the solid pins to find the weaker interfaces at the seams. The pin locations for the latter two samples are highlighted in Figure 8, along with a schematic of the observed crack front progression in the z-direction.

The pin locations for the latter two samples are highlighted in Figure 8, along with a schematic of the observed crack front progression in the z-direction. It appears that the crack front crossed the pin locations at the seams between pins rather than fracturing the pin or pulling it out of the structure. Close inspection of Figure 8a shows a “string” of material originating from the red pin locations and stretching out in a line (left-to-right) toward the next pin location. This is an artifact of a non-optimized pinning process, where the extruded pin has slightly overfilled the hole and excess material has been dragged along the surface to the next pin location. A similar pattern can be seen in the lower row of blue pins in Figure 8a, but this time apparently originating on the right-most pin and being dragged to the left.1774No Pin w/ Pause8-4 Pins, Standard8-4 Pins, StaggeredFigure 7.

Representative fracture surfaces of tensile samples (CF-PLA).A “stair-stepping” effect was observed at the fracture surface, but it is not readily apparent in the photographs in Figures 8a and 8c. The seams in the Standard pin configuration remained constant within a given row. Near bottom of the sample in Figure 8a, each of pins colored blue have a seam at the same layer height, which is 4 layers below the seam locations of the middle row of pins colored red. Observing the fractured sample from the side (schematically illustrated in Figure 8b) showed that the crack front distinctly shifted across the width of the sample to cross at the new seam location. As the crack progressed to the third row (blue pins), the fracture surface shifted up a similar amount to cross the seams of the two pins on the left, but actually shifted back down to the original layer height to cross the lower seam of the pin in the back right (dashed line in Figure 8b).

A similar stair-stepping effect was observed in the Staggered sample in Figure 8c. In this configuration, the seam locations are at the same height for pins that are diagonally positioned. So rather than progressing across rows, the stair-step path progressed diagonally across the sample, continuing to gain height when moving from left-to-right (positive x-direction) in Figures 8c and 8d. The color-coded sections in Figure 8d illustrate the height of the fracture surface that is aligned with the seam of a given pin. These fracture patterns indicate that the weakest location under the current printing conditions is the seam between adjacent pins, which could result from under-filling of voids. Future work will investigate alternate configurations and optimize print conditions.(a){c}(b}y ►I I I I I IFigure 8.

Future work will investigate alternate configurations and optimize print conditions.(a){c}(b}y ►I I I I I IFigure 8. Fracture surface analysis relative to pin position.1775Conclusions and Future WorkThe z-pinning approach has been demonstrated for MatEx samples with a solid cross section. The strength of solid samples without pins was found to be very sensitive to the layer time; less than a one minute delay on every fourth layer caused the UTSZ to decrease by more than 60%. The addition of z-pins to the structure was able to increase the z-strength by ~40% regardless of the pinning configuration used. Inspection of the fracture surface of pinned samples revealed that the crack front navigated to the seams between adjacent pins, indicating that defects may be present due to incomplete filling of the void intended for the pin.

Because solid samples do not have an open architecture, the placement and extrusion of pins is less forgiving than sparse structures and additional optimization of the printing process and pin geometry is required.AcknowledgementsResearch sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Industrial Technologies Program, under contract DE-AC05-00OR22725 with UT-Battelle, LLC.References[1][2][3][4][5][6][7][8][9]Torrado-Perez, A., D. Roberson, and R. Wicker. (2014) "Fracture Surface Analysis of 3d- Printed Tensile Specimens of Novel Abs-Based Materials". Journal of Failure Analysis and Prevention 14 pp. 343-353. Torrado, A.R., C.M. Shemelya, J.D. English, Y. Lin, R.B. Wicker, and D.A. Roberson. (2015) "Characterizing the Effect of Additives to Abs on the Mechanical Property Anisotropy of Specimens Fabricated by Material Extrusion 3d Printing". Additive Manufacturing 6 pp. 16-29. Shaffer, S., K. Yang, J. Vargas, M.

direct Additive Manufacturing of a Cubic Lattice Structure with a Copper Alloy1Jiwon Mun, 1Jaehyung Ju, and 2James Thurman1Department of Mechanical and Energy Engineering, 1155 Union Circle #311098 2College of Visual Arts and Design, 1155 Union Circle #305100 University of North Texas, Denton, Texas 76203-5017AbstractDirect-metal additive manufacturing (AM) processes such as Selective Laser Melting (SLM) and Electron Beam Melting (EBM) methods are being used to fabricate complex metallic cellular structures with a laser or electron beam over a metal powder bed.

Even though their excellent capabilities to fabricate parts with cellular mesostructure, there exist several constraints in the processes and applications; limited selection of materials, high thermal stress by the high local energy source, poor surface finish and anisotropic properties of parts caused by combined effects of one-dimensional (1D) energy based patterning mechanism, the deposition layer thickness, powder size, power and travel speed of laser or electron beam. In addition, manufacturing cost is still high with the Direct-metal AM processes. As an alternative way to manufacture metallic 3D cellular structures, which can overcome the disadvantages of direct-metal AM techniques, polymer AM methods can be combined with metal casting. We may call this “Indirect AM based Casting (I AM casting)”.

We may call this “Indirect AM based Casting (I AM casting)”. The objective of the study is to explore the potential of I AM Casting associated with development of a novel manufacturing process - Indirect 3D Printing based casting which is capable of producing metallic cellular structures within a cell size of 3mm and cell thickness of 0.5mm. We will characterize polymers making sacrificial patterns by PolyJet typed 3D printers; e.g., melting and glass transition temperatures and thermal expansion coefficients. A transient flow and heat-transfer analysis of molten metal through 3D cellular network mold will be conducted. Solidification of molten metal through cellular mold during casting will be simulated with temperature dependent properties of molten metal and mold over a range of running temperature. The volume of fluid (VOF) method will be implemented to simulate the solidification of molten metal together with a user defined function (UDF) of ANSYS/FLUENT.

The volume of fluid (VOF) method will be implemented to simulate the solidification of molten metal together with a user defined function (UDF) of ANSYS/FLUENT. Finally, experimental validation will be followed.IntroductionCellular metals are increasingly receiving attention for their having combinations of mechanical, thermal, and acoustic properties that provide potential opportunities for diverse multifunctional structural implementations. These include ultralight structures with high specific strength [1] and high specific strain [2], excellent impact absorption [3], acoustic insulation [4], heat dissipation media and compact heat exchangers [5].In alignment with recently emerging 3D printing technologies [6], realization of their structural applications of cellular metals appears to become fast.

Direct-metal additive manufacturing (AM) processes which use a laser or electron-beam over a metal power-bed are capable of manufacturing fully complex 3D cellular metallic structures. Selective Laser Melting [7], Electron Beam Melting [8], and Direct-Metal Laser Sintering [9] are the available direct AM processes to fabricate metallic cellular structures. However, in spite of their capability to fabricate parts with cellular geometries, there are still several constraints in the process, which limits their scalability to structural applications. For example, selection of materials is limited: e.g., aluminum alloys are challenging to process due to their high thermal conductivity and high optical reflectivity[10]. Fabricated parts suffer from thermal residual stresses [11]. The 1D energy based patterning mechanism induces poor surface quality mainly by ‘balling’ phenomenon that occurs during laser or electron melting [12].

The 1D energy based patterning mechanism induces poor surface quality mainly by ‘balling’ phenomenon that occurs during laser or electron melting [12]. The ‘balling’ phenomenon limits the direct metal AM process resolution because it causes the formation of discontinuous track, which is responsible for a non-uniform deposition of material on the previous layers, thus inducing a possible porosity and delamination between layers that is detrimental to the functional performance of parts such as fatigue [12]. Direct-metal AM processes also induces anisotropic properties of parts caused by a combined effects of the deposition layer thinness and powder size, power and scanning speed of laser or electron beam [8]. In addition, both costs for machine and manufacturing are extremely high with the Direct-metal AM processes.

In addition, both costs for machine and manufacturing are extremely high with the Direct-metal AM processes. For these reasons, it may be argued that direct-metal AM methods are not capable of fabricating cellular structural parts with a sufficient scale for structural applications.As an alternative way to manufacture 3D cellular metals, which can overcome the disadvantages of direct-metal AM techniques, an AM of polymers can be combined with traditional casting methods. We may call this “Indirect AM based Casting (I AM Casting)”. I AM Casting has been introduced more than two decades ago but quantitative studies on this have not been fully conducted yet. I AM Casting can give more freedom in selecting materials, e.g., aluminum and magnesium cellular structures, with proper design of gate systems related to flow control of molten-metal. It can produce an isotropic property of cell walls.

It can produce an isotropic property of cell walls. It can provide excellent surface quality, which is one of the strong advantages over the direct-metal AM methods. These all provide one with the potential to scale up the cellular metals and other multifunctional structural applications with relatively cheap cost due to the mass-production combined with reduced cost of pattern fabrication using automated 3D printing technologies, which were traditionally used with wax pattern through die casting in investment casting.In spite of its scalable potential of periodic cellular metals to (multi)functional applications, the manufacturing methods of I AM Casting were not fully explored and understood. Only a few qualitative studies have been conducted. A tetrahedral Be-Cu alloy lattice structure, with a cell size of 10mm, was fabricated by investment casting of an ABS sacrificial cellular pattern printed by a Fused Deposition Modeling (FDM) method [13].

Recently, a research group reported a sand casting of cellular structures, whose cell size is about 25mm x 50mm [14] and significant casting porosity was reported. This appears to be associated with network structures with sharp corners of cellular structural molds that may prevent fluid-flow and undesired solidification, resulting in mis- run and porosity.In this study, we explore a new I AM Casting method to build a cellular structure by combining 3D printing of cellular patterns by a polyJet method and centrifugal casting. A transient flow and heat-transfer analysis of molten copper alloy through 3D cellular network mold will be conducted with a sprue system. Solidification of molten metal through cellular mold during casting will be simulated with temperature dependent properties of molten metal and mold over a range of running temperature.

Solidification of molten metal through cellular mold during casting will be simulated with temperature dependent properties of molten metal and mold over a range of running temperature. The volume of fluid (VOF) method will be implemented to simulate the solidification of molten metal together with a user defined function (UDF) of ANSYS/FLUENT, which will be experimentally compared.Preparation of Sacrificial pattern with Additive ManufacturingIn this section, we present a method to produce a 3D cellular pattern fabricated by a 3D printer. A polyjet type professional 3D printer (ProJet™ HD 3500 Plus, 3D Systems) was used to produce a cellular sacrificial pattern. The base material used in the printing is VisiJet® Procast, a castable plastic.

The base material used in the printing is VisiJet® Procast, a castable plastic. The supporting material used for 3D printing is VisiJet® S300, a wax material forhands-free melt away supports.2530) a P M(s s e r t S201510)W m(w o l F t a e H25201510𝑇𝑔 = 97.59 oC𝑇𝑚 = 337.11 oC5500010020030040000.05Strain0.10.15Temperature (oC)Figure 1: Stress-strain curve of VisiJet® Procast under a uniaxial load at a strain rate of 0.001Figure 2: DSC curve of VisiJet® ProcastMechanical Properties: A 3D printed dumbbell-shaped sample is prepared according to ASTM D638 with an extra high resolution printing mode (16 resolution and 750x750x1600 DPI in the x, y, and z directions).

A uniaxial test using a universal testing machine (AGS-X Series, Shimadzu) with a load cell of 5kN and a strain rate of 0.001/s shows a modulus of 549MPa and a yield strength of 8.7MPa (Figure 1), which is about 2-20 times higher than waxes in modulus and strength. It should be noted that waxes have been generally used as materials for building sacrificial patterns in investment casting [15]. The relatively high modulus of the base material of a 3D printer is preferred to the conventional wax materials for fabricating 3D cellular sacrificial patterns with thin walls because the patterns need to bear loads by ceramic slurry or plaster coating.Thermal analysis: For planning burning out of sacrificial patterns, estimation of melting temperature is necessary. We conducted thermal analysis of VisiJet® Procast with a Differential Scanning Calorimeter (DSC). It shows a melting temperature of (Figure 2).

We conducted thermal analysis of VisiJet® Procast with a Differential Scanning Calorimeter (DSC). It shows a melting temperature of (Figure 2). For a reference value, a wax material’s melting temperature was also measured ( ) with DSC as well. In general, waxes have been popular as a material for fabricating sacrificial patterns due to its low melting temperature so that one can easily melt away the wax patterns during the burning out processing to make ceramic mold. The relatively high melting temperature of VisiJet® Procast is not favorable but it can be manageable.Thermal expansion coefficient of VisiJet® Procast was measured with a thermo mechanical analyzer (TMA4000, PerkinElmer); ⁄ which is 140 times lower than that of wax ( ⁄ ). This shows that VisiJet® Procast has the potential to avoid thermal stress related cracking which is generally observed with a wax pattern during burning out process.

This shows that VisiJet® Procast has the potential to avoid thermal stress related cracking which is generally observed with a wax pattern during burning out process. Even considering its relatively high melting temperature of , VisiJet® Procast has about 30 times lower thermal strain than that of wax materials. The density of VisiJet® Procast is ⁄, which is 30% higher than that of wax ( ⁄).500Figure 3: CAD model of a sacrificial pattern of a 3D cellular structureFigure 4: A cellular structural pattern fabricated by a polyjet type 3D printer with VisiJet® ProcastCAD drawing and printing: A complex 3D network cellular structure is pursued to cast with pattern fabricated by additive manufacturing. A computer aided design (CAD) model for the sacrificial pattern of a 3D cellular solid is generated with commercial CAD software, Pro-E, as shown in Figure 3. All necessary dimensions of the 3D cellular pattern are indicated in Figure 3.

All necessary dimensions of the 3D cellular pattern are indicated in Figure 3. Figure 4 shows a cellular structural pattern with internal channels manufactured by a polyjet type 3D printer (ProJet™ HD 3500 Plus) with VisiJet® Procast. The average thickness of the 3D sacrificial pattern is 1.016mm and the standard deviation is 0.027mm.In terms of its high modulus and low thermal expansion coefficient, VisiJet® Procast has the potential to be used as a base material for fabricating cellular structural sacrificial patterns over conventional wax materials.

Especially, the high modulus is good for building cellular structural patterns with thin cell walls and the low thermal expansion coefficient is favorable for lowering thermal stress during the burning out process of the sacrificial pattern.Preparation of a Plaster Mold SystemAfter building cellular patterns from additive manufacturing, a mold for an assembled H2O) (Satin Castsacrificial lattice pattern with a sprue system is prepared. Gypsum (CaSO4 20™, FindingKing Kerr) is used for mold preparation of 3D cellular structural plaster mold.Figure 5 shows the procedure of the preparation of the mold. First, Satin Cast 20TM was weighed for the plaster mold as shown in Figure 5 (1). Then, the quantity of water for mixture with investment was determined, which is recommended to add 40 ml water to 100 g investment powder as shown in Figure 5 (2). Water of was added to investment powder in this experiment as shown in Figure 5 (3).

Water of was added to investment powder in this experiment as shown in Figure 5 (3). At the fourth stage as shown in Figure 5 (4), water and investment powder was mixed with water. Adding and mixing investment powder to water stage through stage (3) to (4) were conducted for 3.5 minutes by recommended work time. After mixing water and investment powder together, this mixture was placed in a vacuum chamber for 2.5 minutes in order to remove the remained air bubbles inside of the mixture as shown in Figure 5 (5). Then, the sacrificial pattern is prepared as shown in Figure 5 (6) and it is covered with a flask as shown in Figure 5 (7). The investment mixture in the vacuum chamber is taken out and is poured into the flask as shown in Figure 5 (8). Then, in order to get rid of the air bubbles inside the flask, the flask with mixture was vacuumed again in the vacuum chamber for 1.5 minutes as shown in Figure5 (9). Then, it is placed outside for two hours to be solidified (Figure 5 (10)).

Then, it is placed outside for two hours to be solidified (Figure 5 (10)). For the burning out process; the furnace was preheated at a temperature of as shown in Figure 5 (11). The flask was placed into the furnace at a temperature of for 2 hours as shown in Figure 5 (13) after removing sprue base as shown in Figure 5 (12). The burning out procedure is as follows; for 2 hours for 2 hours for 2 hours for 4 hours and for 2 hours.Figure 5: Procedure for preparation of a plaster moldAfter mixing VisiJet® Procast is burnt out and only a gypsum plaster mold cavity is left. Table 1 shows thermal and mechanical properties of the gypsum plaster mold.

Table 1 shows thermal and mechanical properties of the gypsum plaster mold. Those properties will be used for simulation in Section 4.Table 1: Material properties of gypsum plaster mold Property Density at Thermal conductivity at Specific heat [31] Roughness Thermal expansion coefficient Preheated temperature Heat diffusivity [31]Value (unit) 650 0.47 1 [ ]Centrifugal CastingExperimentalOnce the cellular structural plaster mold is prepared, molten metal is prepared for casting. The chemical composition and properties of the molten copper alloy (Jewelry Bronze) are shown in Tables 2 and 3, respectively [16].

The chemical composition and properties of the molten copper alloy (Jewelry Bronze) are shown in Tables 2 and 3, respectively [16]. Cu is the dominant chemical composition of the material.Table 2: Chemical composition of the bronze tested in this study [16]MaterialCooper alloy (Jewelry bronze)Min./Max.Unit Chemical Composition (%max, unless shown as range or mean) Cu 91.9Si 4.0Zn 4.0Mg 0.25 MaxPb 0.25 MaxTable 3: Thermo-physical properties of the copper alloy [16] Unit 1035 [ 1005 8774.53 380 173.07 0.00031 18.54Property Melting point – liquidus temperature at Melting point – solidus temperature at Density at Specific heat Thermal conductivity at Viscosity Thermal expansion coefficient atIn this experimental procedure, a centrifugal casting machine (Rey Motorized Centrifugal Casting Machine Rey Industries Inc.) was used for the metal casting as shown in Figure 6.

Figure 6 shows a schematic of the centrifugal casting machine which has a width of 70cm, a length of 70cm, and a height of 90cm. It is composed of eight parts of components. As shown in Figure 6, the casting arm, ③, which has a diameter of 30cm is rotated with the main shaft, ①, as the center surrounded with the base, ②. The experimental flask sample is installed into the flask cradle, ⑤, held by the flask cradle holding arms, ⑥. This flask holder accommodates flasks up to . Then, the crucible holder, ⑦, accommodating up to 30oz is held by the crucible holding arms to melt the casting metal.The copper alloy is melted in the crucible holder with an oxygen-acetylene torch by applying a maximum temperature of 1200 as shown in Figure 7 (a). Meanwhile, flask is preheated up to 482 from an oven then is placed in the flask cradles as shown in Figure 7(b).

Meanwhile, flask is preheated up to 482 from an oven then is placed in the flask cradles as shown in Figure 7(b). Once the flask is installed into the flask cradle, ⑤, the arm mechanism is appropriately balanced with the arm weight adjustment system, ④. Then, the casting arm, ③, is rotated clockwise with a rotating speed of 500rpm for about two minutes.Figure 6: A schematic of the centrifugal casting machineOnce the rotating procedure is done, the plaster flask is removed from the flask cradles to be cooled down at room temperature for about five minutes as shown in Figure 7(c). Then, the removed flask is dipped into water at room temperature in order to quench the plaster mold and casted product as shown in Figure 7 (d). A lattice structure is obtained by cracking the plaster mold and by polishing the final product with a sand blaster.Figure 7: Centrifugal casting procedureThe final product of the lattice structure is shown in Figure 8.

As shown in Figure 8 (a), (b), and (c), very few defects such as mis-run, metal penetration, and air gap were detected and the molten metal filled the plaster mold cavity well. It is known that copper alloys fill the mold cavity well during centrifugal casting due to its high density and low viscosity [17]. On the other hand, it is known that aluminum and magnesium alloys with low density and high viscosity are difficult to fill the mold cavity during centrifugal casting [17]. The thickness of the final product is 1.012mm and the standard deviation is 0.077mm.(a)(b)(c)Figure 8: Final product and defectsThe final product by the centrifugal casting shows a smooth surface.

Direct additive manufacturing methods such as selective laser melting (SLM) and electron beam melting (EBM) are known to have poor surface quality in the parts due to the one dimensional (1D) energy based patterning mechanism combined with the deposition layer thickness, powder size, power and travel speed of laser or electron beam [12]. Due to the poor surface quality, SLM and EBM need a post- process for surface polishing. On the other hand, the indirect AM based casting method shows a promising surface finish.Modeling and Simulation of Centrifugal CastingHydrodynamics of the molten metal flow in the lattice structureFlow of molten metal into the cellular structural mold is estimated using fluid simulation. An inlet velocity into the mold is estimated from the radial speed of the centrifugal machine.

An inlet velocity into the mold is estimated from the radial speed of the centrifugal machine. Considering the angular speed of 500rpm of the rotating arm, the radial speed of fluid particle is obtained.(1) where, is the velocity of the molten metal in the radial direction, is the angular velocity of the rotating arm in ⁄ , is the radius of the casting arm and is the released distance of the weight from the origin of the casting arm. The radial velocity , into the cellular plaster mold and sprue system is obtained as from Equation (1).= √Figure 9 shows dimensions of the lattice structure with holes including sprue system.

The radial velocity , into the cellular plaster mold and sprue system is obtained as from Equation (1).= √Figure 9 shows dimensions of the lattice structure with holes including sprue system. Considering the inlet velocity of ⁄ at the sprue (1 in Figure 10) obtained from Equation (1), the inlet velocity into the cellular mold (5 in Figure 10) is obtained.Figure 9: A schematic of lattice structure and its dimensionsFigure 10: A schematic of the lattice structure with a sprue system to estimate flow speed of molten metalSimulating flow of molten metal in the cluster – cellular mold and strue was carried out. First, the flow type at the inlet of the sprue needs to be identified to check whether or not it is laminar and turbulent at a macroscopic level.

First, the flow type at the inlet of the sprue needs to be identified to check whether or not it is laminar and turbulent at a macroscopic level. The Bernoulli equation with losses is considered to estimate the flow of molten metal after pouring at the sprue gate:𝑔𝑔=𝑔𝑔(∑ )𝑔(2)where, is the height between points of interest, is the pressure, is the density of the molten metal, is the gravitational acceleration, is the velocity of the molten metal, is the frictional coefficient associated with viscosity of metal, is the length of pipe, is the diameter of the channel, and is the loss coefficient at the bend or exit. It should be noted that the energy loss by interface friction associated with resistance of the mold walls to the passage of metal is not considered.Then, the flow speed of the copper alloy at each channel can be estimated using the modified Bernoulli’s theorem with losses in Equation (2).

Table 4 shows initial parameters of the molten metal for estimation of flow regime in the sprue. Table 5 shows the losses used in the sprue system.Table 3: Losses used in the lattice structure in Figure 10SectionRelative roughnessFriction Coefficient (Moody chart)Minor losses coefficient1-2 2-33-4Sudden contraction + entry = = No minor losses = Sudden contraction + entry loss = =The flow speed of the molten metal at each channel can be estimated using the modified Bernoulli’s equation with geometric loss values in Table 5. The corresponding Reynolds numbers were obtained to determine the flow type at each point [18, 19]. It is found that the flow is turbulent through the entire sprue. Two cases of inlet velocity at the gate of the cellular plaster mold are used; 5 and 0.005 for Case 1 and Case 2, respectively, as obtained from Table 6.

Two cases of inlet velocity at the gate of the cellular plaster mold are used; 5 and 0.005 for Case 1 and Case 2, respectively, as obtained from Table 6. Flow simulations without the sprue system are followed with the different inlet velocities as shown in Figure 10.Critical inlet velocity exists to overcome solidification inside of the mold, which is one of the important parameters to avoid casting defects such as mis-run, metal penetration, and air gaps including gating design, initial melting temperature of the liquid metals, etc.

[17], in which the laminar is used for the flow type of the liquid metal based on the calculation of the Reynolds number ( = .Table 4: Reynolds numbers to identify flow typeSection Velocity Magnitude ( )1-2 2-3 3-4 4-5= = = =5Case 1 =Case 2 =Reynolds number 𝑇 𝑇 𝑇 𝑇 Case 1 𝑇Case 2Estimation of heat transfer in the lattice structure with sprue systemIn order to estimate the type of the heat transfer for the numerical simulation, Prandtl number, Rayleigh number, and Nusselt number were calculated using the initial conditions in Table 7.The Pandtl number can be calculated using Equation (3) [20].==𝑚=(3)where, is the kinematic viscosity ( = , is the thermal diffusivity ( =) , is thedynamic viscosity, is the thermal conductivity, is the specific heat, and is the density.

From Equation (3), we obtain a Prandtl number of for Case 1 and for Case 2, which indicates that conduction is dominant compared with the heat convection as shown in Table 8 for Case 1 and Case 2.The Rayleigh number is calculated from the relation with Prandtl number [20];𝑔𝑇 𝑇 = (4) where, is the characteristic length, is the Rayleigh number at position , is the Grashof number at position , is the Prandtl number, is the gravitational acceleration, 𝑇 is the surface temperature (temperature of the wall), 𝑇 is the ambient temperature , is the kinematic viscosity, is the thermal diffusivity, and is the thermal expansion coefficient.

The Rayleigh and number for Case 2, which indicates the conduction flow is dominant and air by natural convection on the mold surface is a laminar flow for both cases, as indicated in Table 8.=obtainedfromEquation(4)isfor Case1The Nusselt number can be calculated as follows [20, 21]:== =(5)where, is the characteristic length, is the thermal conductivity of the fluid, is the convective heat transfer coefficient, is the Reynolds number, is the Prantld number, and is the for heating and for cooling. Here, the casting mold is cooled at room temperature, = is used for this study. The Nusselt number obtained from Equation (5) is for Case 1 and for Case 2, which indicates that the flow type of heat transfer type is conduction for the both cases as well.

Therefore, based on the calculation of the Reynolds, Rayleigh, Prandtl, and Nusselt numbers, the flow type of the molten metal can be considered turbulent and conduction for Case 1 and laminar and conduction for Case 2 are the dominant heat transfer mode in this study.Table 5: Initial conditions for the numerical simulation Initial conditionsValue (unit)Injection velocity (linear velocity)Case 1Case 20Initial temperature of the inlet Flow regimeTurbulentLaminarTable 6: Heat transfer criterion for flow regime justificationCriteriaValueJustification for numerical approachCase 1Case 2Case 1Case 2Prandtl number Rayleigh number Nusselt numberConductionConductionConductionFlow simulation through the lattice structural moldOnce we obtain the initial velocity of molten metal at the gate of the plaster mold for both cases, we go into the estimation of the metal flow in the lattice structural mold.

A commercial finite difference (CFD) code, ANSYS/FLUENT, is utilized to simulate the liquid metal casting to the cellular structural mold [19, 22]. A fully developed turbulent velocity profile for Case 1 and a fully developed laminar velocity profile were applied at the inlet as well as the entire latticestructural mold, which were determined through the criteria of Reynolds number whose velocity input was estimated by the energy equation. An unsteady momentum equation coupled with the unsteady heat transfer (energy equation), the turbulence equations, species equations, and back diffusion equations are applied to solve fluid problems involving solidification of cooper alloy (jewelry bronze) through the cellular structural mold cavity. A Newtonian flow is applied to the simulation under an incompressible condition of the liquid metal.Homogeneous and isotropic properties of molten metal in solid and liquid phases whose properties are shown in Table 3 are applied to the simulation.

A conduction heat transfer is applied to the simulation based on the estimation of heat transfer determined by the Rayleigh, Prandtl, and Nusselt numbers. Surface roughness (2.5 micron) as shown in Table 1 on the wall of the plaster mold is neglected during the simulation for simplifying the problem in our initial attempt to study simulation of molten metal in the cellular structural plaster mold.Figure 11: A schematic of flow simulation of a molten metal in the lattice structural moldCoarse meshes consisting of 1,461,875 elements and 352,218 nodes with a minimum mesh size of and a maximum mesh size of are generated for the flow simulation for both cases. The properties of plaster mold used for the cellular structural plaster mold are shown in Table 1. The constant values of physical properties, e.g.

The properties of plaster mold used for the cellular structural plaster mold are shown in Table 1. The constant values of physical properties, e.g. viscosity, density, thermal conductivity, etc., have been used for the simulation of metal casting in spite of their variation with temperature during casting due to the difficulties in implementing temperature dependent properties in the simulation [23]. This may mislead the solidification, resulting in underestimation of solidification of molten metals which are typically experiencing a temperature varying of ~200C. If the molten metal flows through thin tubes, which is the case of this study with a lattice structural mold, the solidification may be more sensitive to temperature due to more surface area exposed to air.Therefore, a user-defined function (UDF), a used subroutine of ANSYS/FLUENT, where one can implement temperature dependent properties of the molten metal.

Temperature dependent properties of viscosity, density, and thermal conductivity for the molten copper alloy were employed as the input values of the simulation using UDF in ANSYS/FLUENT [19, 24].Continuity and Momentum Equations for fluid flowThe continuity equation for unsteady fluid flow can be written as follows:⃗ = 𝑚 (6) where, is the density of fluid, ⃗ is the velocity vector, and 𝑚 is the mass added to the continuous phase from the dispersed second phase due to the phase change. Here, 𝑚 can be any user-defined source terms according to the problems [19].The conservation of momentum can be described by⃗ ⃗ ⃗ = ̅ ⃗ ⃗ (7)where, is the static pressure, ̅ is the stress tensor, ⃗ is the gravitational body force, and ⃗ is the external body force.

Also, ⃗ can contain other model-dependent source terms such as user- defined sources.Volume of fluid (VOF) method for flow fillTo describe and track the interphase between the solid and the liquid phases, the VOF method was used. This enables to track the transition of the interfaces with arbiter topology and deformation [19, 25]. To define the fractional volume of each phase in the computation domain, the time derivative of is expressed as:= (8)This function has a range from zero to the unity . = indicates that a cell contains no fluid, = indicates the interphase between liquid and solid phases, and = means that a cell is full of fluid.Simulation of solidificationThe fluid flow problems involving solidification taking place at one temperature for pure metals or over a range of temperatures for binary alloys. In this study, an enthalpy-porosity method for the solidification is used instead of tracking the liquid-solid front (VOF method) explicitly [26].

This means that the liquid-solid phase (mush zone) is treated as a porous zone with porosity equal to the liquid fraction [19].The conservation of energy equation can be written as follows [27]:= () (9)where, is the source term, is the sensible enthalpy, is the thermal conductivity, and is the specific heat.The source term derived by Benon and Incropera for a two-phase system can be written as[27]:] (10) where is the thermal conductivity, is the specific heat of the solid phase, is the sensible enthalpy of the solid phase of the solid phase, is the sensible enthalpy, is the mass fraction of the solid phase, is the velocity vector, is the velocity vector of the solid phase, is the sensible enthalpy of the liquid phase, and is the sensible enthalpy of the liquid phase.= [The momentum source in liquid, mush zone (liquid + solid), and solid phases can bewritten as follows, respectively [27].=(11)𝑚 =𝑚 (12)= (13)where, is the dynamic viscosity, is the liquid volume fraction, is a small number (0.001) to prevent division by zero, 𝑚 is a mushy zone constant, is the velocity vector, and is a constant.

The momentum source is proportional to the x, y, and z components of the velocity vector , which has the opposite sign, therefore [19, 27]:= (14)This means that its motion will be reduced to zero (no velocity) as the liquid cools down, implying formation of solid. So, the derivative of Equation (14) with respect to velocity vector becomes a constant or zero. For the temperature-dependent viscosity one needs to assume that liquid becomes more viscous as liquid cools down and its velocity decreases, which implies solidification. The used equations for the temperature-dependent density, viscosity, and thermal conductivity of the copper alloy are shown in Figures 12 through 14.The energy equation for heat transfer is given by𝑇 =(15)where, 𝑇 is the temperature, is the density, is the specific heat, and is the thermal conductivity. In general, viscosity of a material decreases with temperature.

In general, viscosity of a material decreases with temperature. Viscosity of the copper alloy as a function of temperature is plotted in Figure 12.s -m / g K, y t i c o c s i V0.0060 0.0055 0.0050 0.0045 0.0040 0.0035 0.0030 0.0025 0.0020 0.0015 0.001080010001200140016001800Temperature, C◦Figure 12: Temperature-dependent viscosity of molten copper alloy [22, 24]Density decreases with temperature due to the increase in volume. The reason on the rapid decrease between in Figure 13 is due to the phase transformation from liquid phase to solid phase. In the solid phase, the volume which contains atoms is smaller than that of liquid phase.3 ^ m / g k , y t i s n e D8500 8300 8100 7900 7700 7500 7300 7100 6900 6700 650080010001200140016001800Temperatue, C◦Figure 13: Temperature-dependent density of the molten copper alloy [28]Many literatures studied about temperature-dependent thermal conductivity for liquid metal or alloys.

It has been shown that thermal conductivity of liquid metal deceases when its temperature reaches melting point on the assumption that the frequency of vibration of molecules changes near the melting point [29].In many cases, a constant thermal conductivity of molten metal has been used for estimating flow and solidification of molten metal [29].

However, it appears that one needs be careful to implement the thermal conductivity of the molten copper alloy: As can be seen in Figure 14, thermal conductivity is highly nonlinear over a range of 860-980 , which may cause a significant difference in the solidification in molten metal.k - m / W, y t i v i t c u d n o C180 160 140 120 100 80 60 40 20 080010001200140016001800Temerature, C◦Figure 14: Temperature-dependent thermal conductivity of the copper alloy[30]An overall numerical scheme on simulation of molten metal’s flow and solidification coupled with heat transfer is illustrated with a flow chart in Figure 15.

Both flow and heat transfer equations are iterated with updated temperature-dependent properties until the temperature reaches the solidification (1,005 ) of the molten copper alloy.From a preheating temperature of 482 to the solidified temperature of the molten copper alloy, temperature-dependent material properties are updated to both flow and heat transfer equations.Figure 15: A flowchart on the computation of casting using UDF and internal solidification functionThe initial temperature of the copper alloy and the entire plaster mold were maintained as 1200 and 482 , respectively. The time step size used in the simulation was 0.01 . The computational time was about 72 hours with an Intel ® Xeon® CPU E5603 @ 1.60GHz (2 processors) and 24.0 GB RAM.Simulation ResultsSimulation of Flow-fillFlow-fill of molten metal through a cellular structural mold is simulated as a function of time.

Figures 16 and 17 show the volume fraction of the molten copper alloy as a function of time when the temperature dependent material properties for density, viscosity, and thermal conductivity for molten jewelry were applied.Flow directionFigure 16: Simulation of flow-fill at the plane a for Case 1 (Turbulent flow)Figure 16 shows the simulation of flow-fill at a plane along the flow direction for Case 1 with the turbulent velocity flow type. It takes about to fully fill the cellular structural mold with the molten metal with an initial inlet velocity of . Mis-run is detected near the edge (top and bottom) of the cellular structure. The flow-filling pattern shows that the molten metal is first filled in the x direction, followed by filling in the y direction as shown in Figure 16.Figure 17 is the simulation results of flow-fill at the center plane perpendicular to the flow for Case 1. The filling time is the same with that at the xz-plane.

The filling time is the same with that at the xz-plane. Symmetric flow profile isobserved with mis-run at the edge of the lattice mold.Flow direction0 s1.5 s2.0 s Figure 17: Simulation of flow-fill at the xz-plane for Case 1 (Turbulent flow)3.0 sFigures 18 and 19 show the flow-filling of the copper alloy for Case 2 with the laminar flow type as a function of time when the temperature dependent material properties of the molten copper alloy were applied. Figure 18 shows the simulation of flow-fill at a plane along the flow direction for Case 2. It takes about to fully fill the cellular structural mold with the molten copper alloy with an initial inlet velocity of .Flow directionFigure 18: Simulation of flow-fill at the xz-plane for Case 2 (Laminar flow)Figure 19 shows the simulation results of flow-fill at the center plane perpendicular to the flow. It takes about to fill the cellular structural mold with the molten meal with an initial velocity of.

It takes about to fill the cellular structural mold with the molten meal with an initial velocity of . Even though it took about to fill the structural mold at point 5, the casting defect often called mis-run (blue part) was detected. This is due to the slow velocity of the molten copper alloy at the inlet.Flow directionFigure 19: Simulation of flow-fill at the xy-plane for Case 2 (Laminar flow)Solidification (Liquid fraction)Transient solidification of the molten copper alloy through the cellular structural mold is simulated with the temperature-dependent material.

Figures 20 through 24 shows the volume fraction of the copper alloy for Case 1 with the turbulent flow type and Case 2 with laminar flow type as a function of time when the temperature dependent material properties for density, viscosity, and thermal conductivity for molten copper alloy were applied.Figure 20 shows the simulation of transient solidification at the xy-plane along the flow direction for Case 1 with the turbulent velocity profile. The red color represents the liquid phase and blue color means the solidified molten metal. It shows that the molten metal becomes solidified as soon as it passes through the lattice mold.Flow direction .Figure 20: Simulation of transient solidification at the xy-plane a for Case 1 (Turbulent flow)Figure 21 shows the simulation results of the transient solidification at the center plane perpendicular to the flow. Also, the solidification shows similar tendency with that from Figure 20.

Also, the solidification shows similar tendency with that from Figure 20. Also, it can be seen that air gaps, unfilled empty part, at the edge of the cellular mold, were detected.Flow directionFigure 21: Simulation of transient solidification at the xz-plane for Case 1 (Turbulent flow)Figure 22 is the simulation of transient solidification at the xy-plane along the flow direction for Case 2 with the laminar velocity profile. Even though the molten metal was fully filled, many air gaps (red colored parts) were detected as shown in Figure 22.Flow directionFigure 22: Simulation of transient solidification at the xy-plane for Case 2 (Laminar flow)Figure 23 shows the simulation results of the transient solidification at the center plane perpendicular to the flow. Similarly, many air gaps are shown inside of the cellular mold as shown in Figure 23.

Similarly, many air gaps are shown inside of the cellular mold as shown in Figure 23. The solidification along the flow direction is more developed compared to that at perpendicular direction to the flow, which is due to the fact that the velocity magnitude is decreased with the converting of the molten metal flow inside of the cellular mold.Flow directionFigure 23: Simulation of transient solidification at the xz-planefor Case 2 (Laminar flow)Flow-filling time for Case 1 with the turbulent flow regime showed faster filling time compared to that of laminar flow type, as expected. As well as, air gaps at the edges of the cellular structure were detected for Case 1 (turbulent flow). On the other hand, more air gaps were detected which is relatively more spread in overall cellular structure as shown in Figures 22 and 23, which causes degradation in the mechanical properties such as strength, toughness, and etc.

this might due to the fact that the initial inlet velocity of the molten metal is not high enough to flow through thin circuit of the cellular structure and to overcome parameter affecting the fluidity such as surface roughness, friction, and etc.In order to systematically analyze the solidification, liquid fraction (the degree of solidification) of the five points are monitored with the flow-filling time during the simulation time for the both of Case 1(turbulent flow) and Case 2 (laminar flow), as shown in Figures 24 through 29.Figures 24 and 25 show the flow-filling time for both Case 1 (turbulent flow) and Case 2 (laminar flow) at the local points of interest, respectively. Filing time and the solidification time at the local points of interest can be obtained from the flow-fill simulation. Figure 24 shows the filling time from the local points ① to ⑤ as shown on the left in Figure 24, respectively.

Figure 24 shows the filling time from the local points ① to ⑤ as shown on the left in Figure 24, respectively. From Figure 24, it is shown that flow-filling time of the molten copper alloy in the cellular structural plaster mold is about 2s. Once molten copper alloy passes the points of interest, it cools down and solidification occurs. The flow-filling time at each point of interest for Case 1 with the turbulent flow regime is obtained; 0.08s at point 1, 1.1 s at point 2, 2.01s at point 3, 3.86 at point 4, and 4.31 at point 5. The flow-filling time at each point of interest for Case 2 with the laminar flow regime is obtained; 0.35s at point 1, 5.25s at point 2, 9.2s at point 3, 10.35 at point 4, and 20.1 at point 5. It can be obviously seen that the flow-filling time is faster for Case 1 with the turbulent flow type rather than case 2 with laminar flow regime.

It can be obviously seen that the flow-filling time is faster for Case 1 with the turbulent flow type rather than case 2 with laminar flow regime. This is important parameter affecting on the casting process to reduce the casting defects such as mis-run, metal penetration, air gap, and etc.1)%( n o i c a r fm e m u o Vl0.80.60.40.2Point 1 Point 2 Point 3 Point 4 Point 50012345Flow time (s)Figure 24: Flow-filling time for Case 1 (Volume fraction for turbulent flow)1)%0.8( n o i t c a r e m u o Vl0.60.40.20Point 1 Point 2 Point 3 Point 40510 Flow time (s)152025Figure 25: Flow-filling time for Case 2 (Volume fraction for laminar flow)Figures 26 and 27 show the solidification time for both Case 1 (turbulent flow) and Case 2 (laminar flow) at the local points of interest, respectively.

Also, the local solidification time at each point of interest for Case 1 is also obtained; 4.18s at point 1, 2.98s at point 2, 2.27s at point 3, 0.23 at point 4, and 0.01 at point 5.

for Case 2 with the laminar flow, the local solidification time at each point of interest is; 22.2 s at point 1, 3.95s at point 2, 1.05s at point 3, 0.5s at point 4, and 0.13s at point 5.)%( n o i t c a r f d u q i Li10.80.60.40.2Point 1 Point 2 Point 3 Point 4 Point 502 Flow time (s) Figure 26: Solidification time for Case 1 (Turbulent flow)01345)%( n o i t c a r f d u q i Li10.80.60.40.2Point 1 Point 2 Point 3 Point 4 Point 5010 15 Flow time (s) Figure 27: Solidification time for Case 2 (Laminar flow)052025According to the monitored solidification time at the points of interest, the global and local solidification time at the five points were represented for Case 1 (turbulent flow) and Case 2 (laminar flow), as shown in Figure 28 and 29, respectively. Figure 28 shows the flow-filling time and solidification time on the points of interest for Case 1 (turbulent flow) as a function of time.

Figure 28 shows the flow-filling time and solidification time on the points of interest for Case 1 (turbulent flow) as a function of time. The global and local solidification for the both cases shows similar tendency as shown in Figure 28 and 29. However, as represented Figures 28 and 29, it is shown that faster solidification occurs for Case 2 (laminar flow), in particular at point 2. There is limit in the initial injection velocity of the molten metal for the casting process to overcome faster solidification.

That is, if the initial injection velocity of the molten metal at inlet of the casting mold is not enough faster to go through entire casting mold, the molten metal cannot be filled inside the casting mold causing faster solidification.54) s ( e m Ti32Flow-filling time1Solidfiicatioin time (Golbal)0Solidfiicatioin time (local)123 Point of interest45Figure 28: Flow-filling time and solidification time on the points of interest for Case 1 (Turbulent flow)2520) s ( e m Ti15105Flow-filling time Solidfiicatioin time (Golbal) Solidfiicatioin time (local)0123 Point of interest45Figure 29: Flow-filling time and solidification time on the points of interest for Case 2 (Laminar flow)Concluding RemarksAn indirect additive manufacturing method combining 3D printing technology and centrifugal casting was suggested and was implemented to manufacture a 3D cellular structure with a quantitative study on flow simulation of molten metal.

A 3D cellular structural pattern was built with 3D printing of VisiJet® Procast, followed by molding and centrifugal casting. Centrifugal casting with copper alloy was conducted to fill the cellular structural mold cavity made of a gypsum plaster in the experiment. Filling and solidification during the centrifugal casting process were simulated in the plaster mold over a range of running temperature ( ). Temperature-dependent material properties of molten metal were applied for simulation of filling and solidification using a user defined function (UDF) in ANSYS/FLUENT. The major findings through this study are  Additive manufacturing can be used for building sacrificial patterns of 3D cellular structures for casting.

The major findings through this study are  Additive manufacturing can be used for building sacrificial patterns of 3D cellular structures for casting. VisiJet® Procast has the potential to be used as a base material for manufacturing sacrificial pattern over conventional wax materials due to its high modulus and low thermal expansion coefficient: These properties are especially good for making for building cellular structural patterns with thin cell walls and lowering thermal stress during the burning out process of the sacrificial pattern, respectively. I AM Casting shows an excellent surface quality, which can be used for manufacturing cellular structures.A coupled flow and heat transfer of molten metal successfully simulated flow-fill and solidification and is compared with the experiment.More mis-runs were detected in overall cellular structure for the simulation with the laminar flow regime.

On the other hand, fewer mis-runs were observed with the faster flow-filling timeat the edges of the cellular structure for the simulation with the turbulent flow regime of the velocity profile. It is required to get limitation in the initial injection velocity of the molten metal to avoid faster solidification and to overcome bad fluidity of the molten metal for the casting process.The present study with copper alloy will be extended to other metallic alloys for manufacturing light -weight and multifunctional cellular structures in near future.AcknowledgementsThe study received support the Research Initiation Grant (RIG) of the Vice President for Research and Economic Development at the University of North Texas (UNT). The authors also thank KCIS Co. Ltd for partially supporting this study.References [1] Gibson L.J. and Ashby, M. F., 1997, “Cellular Solids-Structure and properties”, Cambridge University Press, Cambridge, UK.

direct laser sintering of corrugated flow field plates for direct methanol fuel cell applicationsKumaran M Chakravarthy and David L Bourell,Materials Science and Engineering Program,Laboratory of Freeform Fabrication, The University of Texas at Austin, Austin, TX-78712AbstractDirect methanol fuel cells (DMFC) hold distinct advantages over traditional hydrogen-based fuel cells. Their commercialization, however, has been bound by many factors – especially their sub- optimal efficiency. This work aims at enhancing the performance of DMFC through the use corrugated flow field plates. Our objectives are two-fold – one, to increase the power density of DMFC by corrugating flow field plates and two, to introduce Laser sintering (LS) as an efficient and robust method for the manufacture of such plates. Corrugated flow field plates with 10% more surface area as compared to a planar design were made by LS & tested in a DMFC environment.

Corrugated flow field plates with 10% more surface area as compared to a planar design were made by LS & tested in a DMFC environment. Our results show that the particle size of the material used – Graphite – has a significant effect upon the green strength of LS parts. We also report the performance of corrugated flow field plates with 10% higher surface area (as compared to planar plates), channel width and depth of 2mm and an electrode area of 5 cm2. This study is the first experimental approach to the use of indirect LS for making such fuel cell components.1. Introduction1.1 Non planar designs in fuel cellsDirect methanol fuel cells (DMFC) use liquid methanol as a fuel, which has higher energy density than hydrogen. In addition, methanol is cheap, easy to store and handle compared to hydrogen. All these present a unique advantage of DMFC over hydrogen based fuel cells.

In addition, methanol is cheap, easy to store and handle compared to hydrogen. All these present a unique advantage of DMFC over hydrogen based fuel cells. However commercialization of DMFC has been slow due to various issues including cost, complexity, low operating voltage and efficiency [1, 2]. Attempts are made to improve the performance of these DMFC [3]. One way of improving the performance of the fuel cells is by improving the power density. Power density of a fuel cell can be increased by increasing the surface area/volume ratio of the fuel cell. Increasing the surface area increases the reaction surface where methanol oxidation and or oxygen reduction takes place thereby enhances the perfogenrmance of the fuel cell. Merida et. al [4] proposed a proton exchange membrane fuel cell [PEMFC] design with a non planar membrane electrode assembly (MEA) architecture that improved the volumetric power density.

Merida et. al [4] proposed a proton exchange membrane fuel cell [PEMFC] design with a non planar membrane electrode assembly (MEA) architecture that improved the volumetric power density. The waved tube cell (WTC) concept proposed has been prototyped with two designs- one where the waved MEA was supported by a ribbed graphite plates and the other with a waved MEA supported by waved expanded metallic structure. Significant gains in volumetric power density have been achieved using the WTC architecture. Li et. al [5] developed and tested a non planar- wave like PEMFC with undulated MEA and perforated bipolar plates. Gains in volumetric power density and specific power with the novel wave like architecture stacks were reported. Laam Tse [6] attempted to improve the power density of the DMFC by corrugating the membrane electrode assembly by pressing and folding methods.

Laam Tse [6] attempted to improve the power density of the DMFC by corrugating the membrane electrode assembly by pressing and folding methods. An analytical model was developed to analyze the relationship between area-to-volume ratio (A/V) and the corrugation aspect ratios. The model predicted a greater increase in A/V for3-D corrugation. Experiments were done to validate the model and it was found that the overall power output was reduced. However since the projected surface area was reduced, the projected power density was increased for MEA corrugations with aspect ratios of 1.25. For MEA with aspect ratios 0.75, the performance was either identical or lower than the projected planar MEA. This performance degradation was attributed to the mechanical degradation introduced during the corrugation process. All the above work has either metallic current collector and/or ribbed graphite plates.

All the above work has either metallic current collector and/or ribbed graphite plates. Metallic bipolar plates have been known to corrode under the conditions seen in a fuel cell.Bipolar plates made from graphite are highly desirable in fuel cell applications due to the properties of high electrical conductivity, low weight and resistance to corrosion that graphite possesses. However, the poor mechanical properties of natural graphite lead to prohibitive machining costs [7]. Indirect laser sintering (LS) provides an alternative way to manufacture bipolar plates using graphite-phenolic system [8]. The main advantage of the LS process is the design flexibility. Corrugated flow field plates, which are difficult to machine into a graphite block, may be made using the LS process and tested for their performance in a DMFC condition.1.2 Green strength of laser sintered partsThe green strength of the graphite-phenolic composites made by LS is not good, 1-2 MPa [9].

Often parts break while trying to remove them from LS machine. This issue becomes complicated when using corrugated plates. So to improve the green strength, attempts were made by adding Nylon/11 [10] and carbon fiber [11]. Nylon/11 additions degraded the green strength and carbon fiber additions improved the green strength. The improvement in strength was offset by a dramatic drop in the electrical conductivity compared to a graphite-phenolic parts.It becomes necessary to reassess the additions and their concentration. In addition, LS process parameters are optimized to observe if improvements in green strength can be obtained without adding new material to the graphite-phenolic system. The effect of graphite particle size and effect of fast cure phenolic resin on the green strength are explored.2.

The effect of graphite particle size and effect of fast cure phenolic resin on the green strength are explored.2. Experimental2.1 Laser sintering and post processing of corrugated flow field platesCorrugated flow field plates with 10% higher surface area than flat plates were designed using SolidworksTM. The channel width and depth were 2 mm and the electrode area was 5 cm2. These plates were made in a DTM Sinterstation 2000 (™) using graphite (#4012, Asbury Graphite Mills, Inc)-phenolic resin (Durite AD 332A, Hexion inc) mix. The parameters used for laser sintering were as follows: fill laser power (22-24 W), outline laser power (4 W), laser scan spacing (0.0762 mm), powder layer thickness (0.1016 mm), partbed temperature (85(cid:176) C) and feed bin temperature (40(cid:176) C). The green plates were carbonized in a vacuum furnace at 1200ºC for 1 hour to convert the phenolic resin to amorphous carbon (brown part).

The green plates were carbonized in a vacuum furnace at 1200ºC for 1 hour to convert the phenolic resin to amorphous carbon (brown part). The brown parts were then infiltrated with cyanoacrylate resin to make the plates fluid impermeable. The infiltrated plates were polished to remove the excess surface cyanoacrylate resin. Corrugated flow field plates with 10% more surface area and 2 mm channel dimensions made by this process are shown in Figure 1 below:Figure 1: Corrugated plates with 10% more surface area compared to a flat plate. The channel dimensions are 2 mm.2.2 Fuel cell testingThe electrochemical performance of the DMFC with the corrugated flow field plates were tested in a single cell fixture with 850 e fuel cell test station supplied by Scribner associates inc. MEA used was a commercial one purchased from fuel cell store [membrane: Nafion™ 117, anode catalyst: Pt/Ru/C, cathode catalyst: Pt/C (4 mg/cm2)].

MEA used was a commercial one purchased from fuel cell store [membrane: Nafion™ 117, anode catalyst: Pt/Ru/C, cathode catalyst: Pt/C (4 mg/cm2)]. The performance of the plates was characterized as a function of clamping torque (65, 70, and 78 lb-in torque), reactant flow direction (counter and parallel flow) and reactant flow rate (oxygen: 0.2 and 0.3 L/min, methanol: 9, 10, and 13 mL/min) to identify optimum parameters from the corrugated plates.For studying the effect of clamping torque, the reactant flow direction was counter flow, methanol flow rate was 9 mL/min, oxygen flow rate was 0.2 L/min. For studying the effect of methanol flow rate, the conditions for testing were as follows: clamping torque: 70 lb-in, reactant flow direction: counter flow; 1 M methanol, oxygen flow rate: 0.2 L/min. For studying the effect of oxygen flow rate, the test conditions were as follows: clamping torque: 70 lb-in, reactant flow direction: counter flow; 1 M methanol, methanol flow rate: 9 mL/min.

For studying the effect of oxygen flow rate, the test conditions were as follows: clamping torque: 70 lb-in, reactant flow direction: counter flow; 1 M methanol, methanol flow rate: 9 mL/min. Fuel cell temperature was kept constant at 65(cid:176) C for all test conditions.Commercial plates with 1 mm channel dimensions were tested under the following conditions: clamping torque: 65 lb-in, reactant flow direction: counter flow; 1 M methanol, methanol flow rate: 6 ml/min; oxygen flow rate: 0.2 L/min, fuel cell temperature: 65(cid:176) C.2.3 Laser sintering process optimization for improved green strength2.3.1 Effect of graphite particle sizeTwo types of graphite particle sizes GS 150 E (~ 75 µ) and GS 75 E (~ 27 µ) obtained from Graftech international were mixed independently with phenolic resin milled to a particle size of ~ 15 µ (novalac based, Durite AD 332 A) obtained from Hexion inc.

The mixes were then loaded into a laser sintering machine DTM Sinterstation 2000 and three point bend specimens (5 in X 0.5 in X 0.125 in) were made with different fill laser powers. The specimens were tested using a5 kN Instron machine (model 3345). The span length was set at 50 mm and a displacement rate of 0.1 mm/min was used. Maximum flexural stress was reported as the flexural strength.2.3.2 Effect of fast curing phenolic resinPhenolic resin contains two parts- a novalac resin and a curing agent- hexa methylene tetra amine (HMTA). Two types of phenolic resins were used in this experiment- one supplied by Plenco as a fast cure resin and other was supplied by Hexion inc- Durite AD 332 A. The Durite resin has ~ 8% of HMTA. Novolac is yellow in color and HMTA is white in color. 8% HMTA Durite is yellow in color and is referred to “yellow” Durite in this work. The new fast cure resin procured from Plenco is white in color indicating higher amounts of HMTA in the mix.

8% HMTA Durite is yellow in color and is referred to “yellow” Durite in this work. The new fast cure resin procured from Plenco is white in color indicating higher amounts of HMTA in the mix. In addition, the fast cure resin has higher amounts of free phenol and water molecules which improve the cross linking properties. Another Durite obtained from Hexion has higher HMTA and is white in color. This resin is called “white” Durite.GS 75 E (27µ) graphite was mixed with “yellow” Durite and Plenco fast cure phenolic separately in a 70:30 ratio by weight, and three point bend specimens were made using the parameters described earlier.

For the Plenco mix, the partbed (70(cid:176) C) and feed bin temperatures (30(cid:176) C) were reduced to prevent the powder mix from caking.The “white” Durite and “yellow” Durite were mixed separately with Asbury Carbon graphite (# 4012) with a particle size of ~75µ in a 65:35 ratio by weight and LS three point bend specimens were made using the parameters described before.3. Results and Discussion3.1 Fuel cell performance tests 3.1.1 Effect of clamping torque The polarization (potential vs. current density) curve and the power density curve for different clamping torque are given in the Figures 2 and 3 below. The performance of the fuel cell with corrugated plates was inferior compared to a fuel cell with commercial plates. The limiting current density for a corrugated plate fuel cell was~150 mA/cm2, whereas the limiting current density of a fuel cell with commercial plate was >300 mA/cm2. This is assumed to be due to the channel size effect.

This is assumed to be due to the channel size effect. The commercial plates had a channel width of 1 mm and the number of passes was ~15 for a 5 cm2 area, whereas the 2 mm channel corrugated plates had 7 passes for the reactants to flow through. This decreased electrochemical area might have offset any potential gain from corrugating the plates. The performance of the corrugated plate will be compared with a flat base line plate of same channel dimensions.The 70 lb-in torque gave better performance than 65 and 78 lb-in torque at high current densities. For low torques, the contact resistance was higher and for higher clamping torques, the gas diffusion layer protruded into the channels, thereby blocking the flow of reactants through the channels. Radhakrishnan and Haridoss [12] studied the difference in structure and properties of different diffusion media and their effect on performance of PEMFC flow field design.

Radhakrishnan and Haridoss [12] studied the difference in structure and properties of different diffusion media and their effect on performance of PEMFC flow field design. They found carbon fiber cloth as a diffusion layer lacked rigidity compared to a carbon paper. Carbon fiber cloth exhibited 43-125% more intrusion compared to carbon paper media. The pressure drop for a 3 mm wide channel was also more with a GDL compared to a 1 mm and 2 mm wide channels. This increased pressure drop may be another reason for reduced performance.Figure 2: Effect of clamping torque on performance of DMFC with corrugated flow field plate (Figure 1), with a 7 pass, 2mm wide channel.Figure 3: Effect of clamping torque on power density of DMFC with corrugated flow field plates (Figure 1).

The performance is slightly better for moderate clamping torque (70 lb-in) at high current densities.3.1.2 Effect of reactant flow direction- counter and parallel flowMethanol solution and oxygen were made to flow in counter current and parallel direction to assess which of these two was favorable for a corrugated plate. From the Figures 4 and 5 below, it can be seen that parallel flow tends to show better performance in the high current region. This is contrary to the normal fuel cells where counter flow gives better performance.Figure 4: Polarization curve for DMFC as a function of reactant flow direction.

This is contrary to the normal fuel cells where counter flow gives better performance.Figure 4: Polarization curve for DMFC as a function of reactant flow direction. The parallel flow tends to give better performance in the high current density region for corrugated bipolar plates.Figure 5: Effect of reactant flow direction: Parallel flow offers better performance than counter current flow for corrugated bipolar plates.3.1.3 Effect of Methanol flow rateMethanol flow rate had a significant effect on the polarization curve of fuel cell with corrugated flow field plates as seen in the Figure 6 and 7 below. Increasing the methanol flow rate from 9 mL/min to 13 mL/min degraded the performance in the high current region. At higher flow rate, the gas diffusion layer could be flooded leading to the poor performance.Figure 6: Effect of methanol flow rate on performance of DMFC with corrugated flow field plates.

At higher flow rate, the gas diffusion layer could be flooded leading to the poor performance.Figure 6: Effect of methanol flow rate on performance of DMFC with corrugated flow field plates. The performance decreases with higher methanol flow rate.Figure 7 shows the power density curve for varying methanol flow rate. Oscillations seen in the curves may be from corrugated channels in the flow field plates. The peak power also shifted to lower current density with increasing flow rate.Figure 7: Effect of methanol flow rate on power density of DMFC with corrugated flow field plates. The power density decreases significantly with increasing the methanol flow rates. There are oscillations in the power density curves with increasing flow rates.3.1.4 Effect of Oxygen flow rate Increasing the oxygen flow rate from 0.2 L/min to 0.3 L/min degraded the performance of the fuel cell in high current region, Figures 8 and 9.

Oscillations seen in both the polarization curve and power density curve could arise from corrugation in the flow channels. Also, the oscillations were more pronounced with higher flow rate, which was similar to the methanol flow rate experiments.Figure 8: Effect of oxygen flow rate on performance of DMFC with corrugated flow field plates. Increasing the flow rate deteriorated slightly the performance in the high current region.Figure 9: Effect of oxygen flow rate on the power density of DMFC with corrugated flow field plates. With a higher oxygen flow rate, the peak power decreased and shifted to left.3.2 Laser sintering process optimization for increased green strength3.2.1 Effect of graphite particle sizeFigure 10: Effect of graphite particle size on green flexural strength of LS specimens. The graphite particle size significantly influenced the green flexural strength.The green flexural strength was strongly influenced by the graphite particle size [Figure 10].

The graphite particle size significantly influenced the green flexural strength.The green flexural strength was strongly influenced by the graphite particle size [Figure 10]. With a higher graphite particle size, the flexural strength was higher. This had been postulated to be due to the number of bridges formed by a finer phenolic resin around a large graphite particle which would be higher than that on a smaller graphite particle [8]. SEM micrographs for small graphite particle specimens [Figure 11a] show that the molten phenolic did not wet the graphite. Instead it re-solidified into fine globules. SEM micrographs of larger graphite specimens [Figure 11b] show that the phenolic resin was able to flow and wet the graphite particles. The BET surface area of 75µ graphite was ~4 m2/g; while the surface area of 27µ graphite was ~12 m2/g.

The BET surface area of 75µ graphite was ~4 m2/g; while the surface area of 27µ graphite was ~12 m2/g. Cross sectional SEM micrograph [Figure 12] shows that interlayer bonding is better for samples with large graphite particle size than samples with small graphite particles. This may be another reason for the difference in strength values. Also, the mix with large graphite particle size responded to laser power better than the mix with fine graphite size.(a)(b)Figure 11: SEM micrographs (planar view) of a sample with 27µ graphite (a) and sample with 75µ (b) graphite. Note the phenolic is seen as small molten globules in (a) and is completely wetting graphite in (b). The BET surface area of the graphite particle is ~ 12 m2/g (a) and ~4 m2/g (b).(a)(b)Figure 12: SEM micrographs (cross sectional view) of a sample with 27µ graphite (a) and sample with 75µ (b) graphite.

Note the presence of large voids in (a) where as good interlayer bonding is seen in (b).3.2.2 Effect of fast cure phenolic resin(i) Plenco fast cure resin vs. yellow DuriteThe Plenco fast cure resin when fed inside the laser sintering machine caked in the feed bins. This was attributed to the melting of free phenol in the resin and to prevent caking, temperature of the bins and partbed were lowered to 30(cid:176) C and 75(cid:176) C respectively. Parts made with those parameters yielded a flexural strength of ~ 0.45 MPa, which was same as that of regular resin[Figure 13]. This low value could also be due to the graphite particle size effect which might have reduced the influence of fast cure resin. SEM micrographs [Figures 14, 15] of samples made with both resin mixes showed similar features.Figure 13: Effect of fast cure resin on green flexural strength of LS specimens.

SEM micrographs [Figures 14, 15] of samples made with both resin mixes showed similar features.Figure 13: Effect of fast cure resin on green flexural strength of LS specimens. Graphite particle size was ~27µ.(a)(b)Figure 14: SEM micrographs (planar view) of a sample with yellow Durite (a) and sample with Plenco fast cure resin (b). The graphite particle size was 27µ graphite. Both the micrographs look similar with non-wetting fine phenolic molten globules. The strength values were also similar (~0.45 MPa).(a)(b)Figure 15: Cross sectional SEM micrographs of a sample with yellow Durite (a) and sample with Plenco fast cure resin (b). The graphite particle size was 27µ graphite. Both the micrographs look similar with large porosity. The strength values were also similar (~0.45 MPa).(ii) White Durite vs. yellow DuriteThe mix with 35 wt% white Durite gave strength levels ~ 3.5 -3.75 MPa.

The strength values were also similar (~0.45 MPa).(ii) White Durite vs. yellow DuriteThe mix with 35 wt% white Durite gave strength levels ~ 3.5 -3.75 MPa. This strength level was comparable to those achieved with up to 25 wt% carbon fiber additions [13]. The fill laser power seemed to have less effect on this composition. All the white Durite specimens made with 20-30 watts gave similar strength values [Figure 16]. The SEM micrograph of a white Durite specimen (Figure 17b) appears to be denser than the yellow Durite specimen (Figure 17a). A higher level of Hexamine in the white Durite resin leads to more densification than the yellow Durite resin. The mix with 35 wt% yellow Durite had slightly lesser strength levels (2.75- 3.5 MPa) and the difference was more pronounced when processed at lower energy levels. Yellow Durite as a resin is responsive to the laser power.

Yellow Durite as a resin is responsive to the laser power. The improvement achieved with white Durite over yellow Durite was comparable to the order of the strength of finer graphite mix (~0.4 MPa).Figure 16: Effect of fast cure phenolic resin on the green flexural strength of LS specimens.(a)(b)Figure 17: SEM micrographs (planar view) of 35wt% yellow Durite (a) and 35wt% white Durite (b) specimens. White Durite specimen is denser than the yellow Durite specimen.The advantage of these mixes over the carbon fiber additions was the electrical conductivity of the brown/finished part would be higher because phenolic would change into conducting amorphous carbon after burnout. The electrical conductivity of specimens with 25 wt% carbon fiber was ~ 80 S/cm [11], which was lesser than the DOE requirement, 100 S/cm [14].4. Conclusions and Future work1.

The electrical conductivity of specimens with 25 wt% carbon fiber was ~ 80 S/cm [11], which was lesser than the DOE requirement, 100 S/cm [14].4. Conclusions and Future work1. A corrugated flow field plate design with 10% more surface area compared to a planar design was made and tested for its performance in a fuel cell. New designs with higher corrugations and channel dimensions comparable to the commercial plates will be made and tested for their performance.2. The green strength of the laser sintered parts was optimized by studying the effect of graphite particle size, fast cure phenolic resin and changing the composition of the mix. Graphite particle size had a significant effect on the green strength and larger particle sizes are preferred for higher green strength.3. Two types of fast cure phenolic resin were tried. One resin (Plenco) was difficult to process inside the laser sintering machine due to caking.

Two types of fast cure phenolic resin were tried. One resin (Plenco) was difficult to process inside the laser sintering machine due to caking. The laser sintering process parameters were modified significantly to enable the resin to be processed. The mix with higher HMTA (white Durite) resin in a 35: 65 (phenolic: graphite) ratio gave strength level comparable to a mix with 25 wt% of carbon fiber addition. AcknowledgementThis work was supported by the Office of Naval Research MURI grant no. N00014-07-1- 0758. The authors thank the support of Mr. Albert V. Tamashausky of Asbury Graphite Mills, Inc. for discussions and help with supplying graphite used in this work. We are also grateful to Dr. Rajan Srinivasacharaya and Mr. William Handel of Hexion specialty Chemicals Inc, Louisville, KY for discussions and for supplying phenolic resin used in this work. We are also grateful to Dr.

tal Compositethe SLS ProcessJamesBadrinarayan2, J. W. Barlow2, J. J. Beaman l , andBourelll ,31. Department of Mechanical Engineering Department of Chemical Engineering Center for Materials Science and Engineering1. Department of Mechanical Engineering Department of Chemical Engineering Center for Materials Science and EngineeringABSTRACTsintering, an intermediate polymer binder is combined with powder to produce green preforms with the Selective Laser Sintering (SLS) process. To produce parts with desirable strength dimensional control, the binder is gradually removed from the green preform (obtained from the SLS process), and the remaining form is lightly bonded. This porous part then infiltrated. Final part density, shrinkage, and strength data are presented. An injection mold insert was fabricated from this material and used to PMMA, polyester, and polycarbonate parts.

Final part density, shrinkage, and strength data are presented. An injection mold insert was fabricated from this material and used to PMMA, polyester, and polycarbonate parts. To date, the mold insert mold and meltAs a near term alternative process to direct176 shots at injection up to 300°C.to 35,000temperaturesinjection molding of thermoplastics are traditionally manufactured from cast metal by precision metal cutting devices. costs are correspondingly high. These high costs and production times effectively preclude preparation of prototype plastic parts by injection molding. Prototype plastic processes.

costs are correspondingly high. These high costs and production times effectively preclude preparation of prototype plastic parts by injection molding. Prototype plastic processes. Unfortunately, parts are consequently usually manufactured by the properties of such parts may not be truly representative of those of injection molded parts due to the thermal stresses and molecular orientation inherent to the injection molding process.Molds toused formachining and polishing times are quite extensive andClearly, there is a great need for low cost, low production time molds which could produce the 50 - 500 plastic parts needed to prove a design. The University ofTexas.at Austin is.presently investigating new material systems and processes that potentially permit· the rapid· fabrication of molds by SLS and subsequent processes.

The University ofTexas.at Austin is.presently investigating new material systems and processes that potentially permit· the rapid· fabrication of molds by SLS and subsequent processes. •In •addition to cost and time savings, potential advantages to mold making by the SLS process include: incorporation of custom cooling channels and other features that are not easily machined, ease of manufacturing duplicate molds, opportunity to evaluate the runner system prior to machining a production mold, and ease of cavity design by use of commercial solid modelers.system and process include: design, and sufficient strength withstanding injection pressures and temperatures. With most powder processes'ithere is usually a trade-offbetweeuJ.ninimal shrinkage and strength. In an attempt. to minimize shrinkage issues. and.

With most powder processes'ithere is usually a trade-offbetweeuJ.ninimal shrinkage and strength. In an attempt. to minimize shrinkage issues. and. still· have enough strength, a powder/infiltrant system was chosen as the first material system for investigation.Two of the important goals of theminitnal, predictable, shrinkage from303DISCUSSIONThe target process for manufacturing molds in the SLS process is outlinedin figure 1:CAD Solid Model of Desired Part (positive)CAD Solid Model of Mold (negative)Add Runner System Design - ( V ) SLS Processed Green Mold n Debind/Bondto Mold Solid ModelInfiltrate Porous MoldInstall Mold and Inject PartsFigure 1. Process for fabricating moldsIn an effort to arrive at this process, the material system and bonding cycleperformance needed to be evaluated. To detennine preferred bonding cycle and material system, green three point bend bars and 11/4" thick blocks were fabricated with different powders.

To detennine preferred bonding cycle and material system, green three point bend bars and 11/4" thick blocks were fabricated with different powders. Different bonding cycles were evaluated to determine shrinkage and strength of the fired bars prior to infiltration. Strength measurements were also taken on three point bend bars that were infiltrated with a total of four different infiltrant systems.An initial preferred powder/infiltrant system was selected. Prior to fabrication of the mold insert using the SLS process, a mold insert was fabricated with the process outlined in figure 2. The insert was fabricated in this manner because the selective laser sintering equipment available at the university is not capable of manufacturing parts of the required scale.

The insert was fabricated in this manner because the selective laser sintering equipment available at the university is not capable of manufacturing parts of the required scale. However, strength tests on bars made both in the SLS process and in the oven show the resultant material systems to be effectively the same.Pour Powder/Binder into Large Aluminum Box..Heat in Oven to Consolidate Green "Slab"I.. i - lRemove Slab from Box H Deblnd/BO~Infiltrate Porous SlabMachine Mold Insert from Slab (CNC)Install Mold Insert and Inject PartsFigure 2. Process for fabricating molds for material evaluation304Figure 3 shows a photograph of a completed mold assembly. For testing purposes, a standard Master Unit Die plaque mold was modified by replacing the steel plaque cavity with the insert. As shown in the foreground, the insert is 8.4" x 5.3" x 0.6" and the part cavity is 5" x 2.5" X 0.25".

As shown in the foreground, the insert is 8.4" x 5.3" x 0.6" and the part cavity is 5" x 2.5" X 0.25". It is equipped with seven standard ejector pins (4 through holes in the part cavity, and 3 through holes in the runner cavity). The runner is 1/4" in diameter and the letters are 1/16" deep. The insert was not channeled for water cooling. For this reason, the cooling times were kept long (near 55 seconds).Figure 3. Mold AssemblyRESULTSTable 1 shows the average density and strength measurements taken on 10SLS processed three point bend bars with the same material used to fabricate the tooling insert.TABLE 1 SLS Processed Three Point Bend Bar DataRelative Density (%)Modulus of Rupture (psi)Green:51421Porous:3295Infiltrated:983,841305These same bars demonstrated linear shrinkages of 0.45% in the X direction, 0.82% in the Y direction, and 1.41% in the Z direction.

To date, the 1 1/4" thick blocks have also shown linear shrinkages of less than 2% in all directions after bonding.Table 2 shows the molding conditions used to inject the fITst 176 shots intothe mold insert.TABLE 2 Molding ConditionsPolymer# of ShotsPeak Inj. Press. (psi)Temperature CC) zones 1,2, 3, & 4Injection Time (sec)ABS1018000200,190,200,2003PMMA2835000175, 170, 160, 1603Kodar A1502235000275,275,275,27542535000245,275,300,3004Additionally, for all plastics, the injection hold pressure time was set to 2 seconds and the cooling time was set to 55 seconds.Observations of the molded parts show only 2 defects during the 176 shotrun. Both of these defects occurred while molding PMMA. At shot 105 a hairline fracture (perhaps surface) started at one of the ejector pins and grew over the next 25 shots to a total length of approximately one inch.

At shot 105 a hairline fracture (perhaps surface) started at one of the ejector pins and grew over the next 25 shots to a total length of approximately one inch. The crack did not noticeably propagate further after that point and did not affect the performance of the mold other than the faint mark left in the moldings. At shot 123, a small but noticeable chip of material around another ejector pin broke loose. This breakage progressed through shot 127 to a total size of approximately 1/16" x 3/16" x 1/16". This defect, similar to the hairline fracture, did not grow further. We are currently in the process of evaluating mold erosion by measuring dimensions of the molded parts.CONCLUSIONSThe generally low shrinkages associated with post processing green shapes suggest acceptably good geometric fidelity for successful molding of plastic parts. Our evaluation of the mold material is presently incomplete, however, these initial results are most encouraging.

Our evaluation of the mold material is presently incomplete, however, these initial results are most encouraging. They suggest that even materials with relatively low strengths can be adequate for prototype insert service, provided reasonable care is exercised in the overall tool design. Mold erosion does not qualitatively appear to be significant. The deficiencies seen near 2 out of 7 ejector pins are of some concern. We are presently developing a higher strength material that could correct this problem. Ejector pin sleeves could also be installed in future designs.We intend to prepare new inserts with cooling channels and variousgeometric features to gain additional understanding of the relationships between insert material properties, mold design, and molding conditions.306ACKNOWLEDGMENTSThe authors gratefully acknowledge the financial and general support of the DTM Corporation, Austin, Texas, for portions of this work. J.

J. Tobin especially wishes to thank Brian Carr of the DTM Corporation and Hank Franklin of the UT Austin Mechanical Engineering Machine Shop for their assistance.307.

direct Rapid Molds for Prototype Lost-Foam Pattern Production V. R. Gervasi*, F. Z. Shaikh** *Rapid Prototyping Center, Milwaukee School of Engineering, Milwaukee, WI 53202 **Scientific Research Lab, Ford Motor Company, Dearborn, MI 48121AbstractLost-foam (also known as Expendable Pattern Casting, EPC) is an ever-growing metalcasting technique, capable of producing complex metal components without parting lines. Mold preparation for lost-foam casting is typically accurate, but expensive and slow. The goal of this research was to develop a new approach for producing rapid lost-foam molds. With this new approach, patterns generated by SFF technology are used to form indirect composite lost-foam molds. Ultimately, our objective is to produce these molds quickly, accurately, and inexpensively.

Ultimately, our objective is to produce these molds quickly, accurately, and inexpensively. This new approach to lost-foam mold-making will be explained as well as the results of one trial.IntroductionLost-foam casting is an increasingly popular method for producing complex metal parts. The process starts by molding simple, low-density, foam segments with minimal undercuts. Next, segments are adhered together, to form elaborate lost-foam patterns (figure 1) --often with intricate features, such as cooling channels and oil passages. Then, foam patterns are attached to a sprue and runners to form a tree, and coated lightly with a refractory coating. Next, assembled trees are placed in a flask and loose, unbonded sand is vibrated around the tree. The top of the sprue is exposed at the surface of the sand. Finally, molten metal is poured directly into the exposed portion of the tree, replacing the tree and its foam patterns, with exceptional accuracy (figure 2).

Finally, molten metal is poured directly into the exposed portion of the tree, replacing the tree and its foam patterns, with exceptional accuracy (figure 2). This process, first used to produce bronze figurines, has advanced considerably over the years. Today, many lost-foam patterns are produced for many applications, including automotive engines, valve bodies, and electric motor housings. Foam patterns are also used for investment casting, replicast [1], and packaging applications.Figure 1. Polystyrene PatternFigure 2. Lost-Foam CastingFoam patterns are produced using a polystyrene-foam injection molding machine andspecialized tooling. The lost-foam injection molding machine consists of several main components including:Bead hopper --from which polystyrene bead is continually supplied to the fill-gun. • Steam chests --stationary steam-chest and movable steam-chest.

• Steam chests --stationary steam-chest and movable steam-chest. • Mold halves --mold halves are mounted on and closed between the two steam chests, to form the mold cavity.506Steam vents & fill guns --molds are equipped with numerous steam vents and several fill- gun holes. Fill-gun tips, attached through the stationary steam-chest, are positioned flush with the surface of the mold cavity, through the access holes.Ports & drains --each steam chest is equipped with a steam inlet port, a compressed air inlet port, a water-cooling sprinkler system, and a drain.The fundamental process used to form a polystyrene pattern can be described using fivemain steps as follows:Step 1. The drains on both steam chests are opened and partially-expanded polystyrene bead iscarried into the mold cavity by a stream of compressed air. Excess air within the mold cavity is vented through steam-venting while the bead is packed within the mold cavity.

Excess air within the mold cavity is vented through steam-venting while the bead is packed within the mold cavity. Steam venting, located in the mold wall, communicates with the lower pressure steam chests and drains.Step 2. Drains are closed and steam (250 F, 15-25 psi) is applied to the bead through the steam- venting. Thermally expanding pentane gas within the bead causes it to expand until the mold cavity is filled. The fine slits on the steam-venting prevent the bead from expanding into the steam chest through the mold wall. The steam venting also allows steam to heat the bead and induce bead expansion and fusion, forming the final polystyrene pattern.Step 3. Steam is turned off and the drains are opened. The water sprinklers spray water on the mold posteriors, cooling them rather quickly. The polystyrene pattern cools through the mold-wall to a temperature where it can be ejected.Step 4.

The water sprinklers spray water on the mold posteriors, cooling them rather quickly. The polystyrene pattern cools through the mold-wall to a temperature where it can be ejected.Step 4. Finally, the moveable steam chest retracts, opening the mold for pattern ejection, and theprocess is repeated.Step 5. Families of pattern segments are joined with adhesive to form complex patterns.This is a simplified description, several common steps such as pre-expansion of bead, pre-heating, steam-through, and vacuum-assisted filling have been omitted in this simplified explanation. The lost-foam molding process does not experience the pressures common with injection-molding, lost-foam molds are designed to handle 50psi.Traditionally, lost-foam tooling is CNC machined from either a billet of aluminum or anear-net casting. Tooling is usually thin-walled (_ to 3/8 inches) to allow steam transfer from the back of the mold to the mold-cavity.

Tooling is usually thin-walled (_ to 3/8 inches) to allow steam transfer from the back of the mold to the mold-cavity. Venting is added by drilling holes normal to the surface of the cavity and steam-vents are press-fit until flush with the surface of the mold. Traditional approaches result in a tool with excellent thermal conductivity, accurate features, at a high cost ($10,000 on the low end) and long lead-time (4 weeks on the low end).Several new processes for producing molds used for lost-foam pattern production havebeen enabled by SFF. These new processes may someday produce molds at a fraction of the cost and time of existing methods, along with a potential improvement in accuracy and complexity. Several direct and indirect SFF based processes have been explored. Direct approaches include507SLS metal, pro-metal, POM, or Lens and indirect approaches include spray metal or composite casting.

Direct approaches include507SLS metal, pro-metal, POM, or Lens and indirect approaches include spray metal or composite casting. When exploring the feasibility of using direct SFF tooling, it was concluded that these processes have much potential in the future, but the materials today are limited. Limitations including high shrink factors, steam-venting problems, high cost, small build envelopes, and low working temperatures (mainly for polymer based SFF) were identified. Indirect methods, such as spray metal and composite casting were also considered, both having advantages and disadvantages. For these two indirect methods, a clay parting line is handcrafted around the positive RP pattern. The first mold-half is spray-metal coated and back-filled with a composite material, followed by clay removal. Next, the second mold-half is spray-metal coated and back- filled. After composite solidification, the pattern is removed and the two mold halves are ready for machining.

Next, the second mold-half is spray-metal coated and back- filled. After composite solidification, the pattern is removed and the two mold halves are ready for machining. Although spray metal tooling has good surface characteristics, it becomes limited when deep features are required. The second approach skips the spray-metal step and casts a composite tool from the RP master. Adding steam venting in both methods becomes difficult and weakens the mold. Steam venting placed after mold construction can loosen, reducing part quality, cycle-time, and tool life. These two indirect approaches result in a solid rather than thin- walled mold, further complicating steam vent addition.Other labor-intensive approaches are used to make thin-walled molds from a positive SFFmaster, but the drive behind this research was to let current technology minimize human intervention.

For this reason attention was focused on combining RTV silicone molding with composite casting to make a mold from an SFF pattern. In addition, a novel cast-in-place steam venting technique was used to reduce machining.ObjectiveThe main objective of this research was to use a new, indirect composite toolingtechnique to produce molds for the lost-foam pattern making process. The molds would be tested under near-typical molding conditions to determine if this technique is worthy of further development, and if so, which steps need refinement.Another objective was to estimate cost and time requirements for this process if it were tobe used in an industrial setting.ApproachThe approach used consists of eight main steps starting with CAD model of mold halves and ended with polystyrene injection molding (mold testing in this case). The part geometry chosen for this process is shown in figure 3.

The part geometry chosen for this process is shown in figure 3. The part is a section of a four-cylinder engine block and has dimensions of 17x15x3 inches. The process used to form this part is described in the following eight steps.CAD Model of Mold HalvesCAD models of both mold halves were designed based on the part geometry. The mold face was approximately 3/8 of an inch thick and supported by a web of ribs. Walls were added around the mold face and ribs.508RP of Mold PatternsLOM was the SFF process selected for this project due to the large build envelope and relative speed. LOM patterns produced were in the form of the final mold rather than the final part.Silicone MoldingLOM patterns were mounted in twomold boxes and RTV silicone was poured around each. Specialized mold boxes, a large molding chamber and a silicone injection system were required to fill the large molds under vacuum.

Specialized mold boxes, a large molding chamber and a silicone injection system were required to fill the large molds under vacuum. After silicone vulcanization, the LOM patterns were removed from the silicone molds.Steam Vent PlacementWith other, indirect mold makingFigure 3. Part Geometryprocesses, steam-venting has a tendency to become loose, and sometimes is driven out by the cyclic bead pressures. For this new approach, Steam-venting was cast-in-place to reduce machining and improve performance. Traditional Steam-vents were modified to prevent the composite slurry from entering the steam passageway and slits. Steam-vent faces were positioned firmly against the silicone surface before silicone mold assembly and composite casting.Composite Mold CastingAfter the silicone halves were assembled within the mold boxes, they formed a void, which embodied the form of the final mold. Steam-vents remained in position, within the void.

Steam-vents remained in position, within the void. The composite material used was a high temperature aluminum filled epoxy. The composite slurry was injected into the silicone molds under vacuum and then room temperature cured at 80 psi. The silicone was stripped from the final molds and the composite molds were trimmed of excess composite.Post CuringThe composite molds were placed on a flat aluminum plate positioned in an oven.The composite molds were buried in sand to reduce thermal shocks and hold the molds flat. Both molds were incrementally heated to above 300 degrees F through a series of 50 degree steps. After cooling, the molds were cleaned and prepared for final machining.Final MachiningMachining consisted of opening the backs of the steam vents and fitting the mold intothe fixture. The fixture was used to hold the tooling halves in place while foam bead was injected, expanded, and cooled.

The fixture was used to hold the tooling halves in place while foam bead was injected, expanded, and cooled. The fixture was designed to take the brunt of the molding press clamping force while holding the mold-halves in alignment -through a 150ºF-temperature range.509The tools expand due to thermal expansion and must be held to expand together. Parting-line alignment is critical to hold accuracy and prevent crashes.Mold TestingAfter the mold halves were fitted into the fixtures and bolted onto the steam-chests ofthe foam injection machine they were handled like a traditional aluminum mold. Autoclave pressures between 15 and 25 psi were used. After part formation the mold was opened slightly and the pattern was blown to the movable side. Next the mold was opened completely and the part was blown off the cavity using air-eject. The main difference when “tweaking” this mold was the initial steam pressures.

Next the mold was opened completely and the part was blown off the cavity using air-eject. The main difference when “tweaking” this mold was the initial steam pressures. After several shots the pressure was increased to 25 psi.Results & DiscussionThe results of this research are best divided into three main categories including: 1) processand testing, 2) process lead time, and 3) composite mold costs.Process & TestingThis new mold-making approach was able to produce a mold for testing, images fromeach step are shown in figure 4. CAD modeling of the mold halves was done using Magics RP by Materialize. This was somewhat challenging due to a combination of CAD geometry and .stl quality. As shown, ribs and outer walls were added to a thin-walled mold face. Twenty three shut-offs were required for this part geometry.A LOM 2030 was used to produce patterns of the mold halves. One pattern was generated and decubed with little to no difficulty.

Twenty three shut-offs were required for this part geometry.A LOM 2030 was used to produce patterns of the mold halves. One pattern was generated and decubed with little to no difficulty. The other mold half was problematic due to bad edges and a shift during the build. The problematic pattern was repaired manually to reduce the undercut, but it was never completely eliminated.The silicone molding, steam vent placement, and composite casting steps were almost problem free. The greatest difficulty was adding the steam-vents and assembling the silicone halves without knocking a vent out of place. Of the 600 steam vents only one vent shifted. The silicone captured and transferred mold features exceptionally (including undercuts). The vent faces were clearly visible on the mold face with little to no composite blockage.Post-curing was a step of concern. The molds had the potential to curl, warp, or distort atthe elevated temperatures but they did not.

The molds had the potential to curl, warp, or distort atthe elevated temperatures but they did not. The machining took a considerable amount of time, mainly do to the generous composite clearance we added behind the steam vents. All 600 vents were opened from the back of the mold by breaking through the 1/8 inch thick composite barrier that remained. Cutting through with a rotary tool was not the difficulty, but finding the location to start cutting was. Vents were securely cast in place. Machining the mold halves to fit in the fixture was straight-forward and required minimal machining.510The focus of mold testingwas on the composite mold integrity and pattern quality. The composite molds were designed to operate in temperatures cycling from 250 deg F down to 60 deg F repeatedly and pressures reaching 60 psi without deflecting or cracking. Ribs were added to support the mold cavity during bead injection and expansion.

Ribs were added to support the mold cavity during bead injection and expansion. The molds withstood the harsh conditions for approximately 10 hours with no notable wear or deflection. Twenty three areas of the mold with shut-offs also functioned well, with little to no flash as shown in figure 5. The fixture used, to hold the mold halves together, performed as expected.CAD Design of ToolRP PatternSilicone MoldCore Vent PlacementCreate Composite MoldPost CureFinal MachiningNone of the patterns produced were of castable quality, although much potential was evident. The mold halves were modified several times during testing to improve bead fusion and fill --resulting in moderate improvements. 68 steam-vents were added to one side of the mold, but there was still not enough bead reaching some regions for proper fill and fusion. Also, undercuts due to the poor LOM pattern made part ejection difficult.

Also, undercuts due to the poor LOM pattern made part ejection difficult. These undercuts were sanded but could not be removed sufficiently for smooth part ejection. Ejector pins were needed to push the patterns past these undercuts. Better LOM patterns, produced on a newer LOM machine, would have reduced or eliminated these undercuts.TestingFigure 4. Mold Making ProcessCast-in-place steam vents and added steam-vents performed well without any dislocation (approximately 600 steam- vents total). Steam transfer is somewhat limited in this steam-vent design. A new design, providing ease of use and better steam delivery is recommended for this rapid tooling process. The cost of preparing steam-vents was expensive and could be reduced with a new design and automated steam-vent production.Figure 5. Pattern with shut-offs visible511Process lead timeFigure 6 illustrates the amount ofTime Expenditurestime spent on each step of the process.

Pattern with shut-offs visible511Process lead timeFigure 6 illustrates the amount ofTime Expenditurestime spent on each step of the process. All steps include training as well as focused work. Portions of the work, RP pattern, silicone molding, composite molding and post curing are unattended activity and continue around the clock. The remainder of the steps would likely take place during a normal work day. Using data from figure 6, figure 7 was created to illustrate the estimated time required for trained lost-foam mold makers to produce polystyrene patterns via this process (given ideal conditions). The time is around two weeks and could be faster if mold halves were produced in parallel using two LOM machines. The time to metal could be reduced if the foundry was included in early planning.12010080s r u o H6040200ActivityFigure 6.Figure 8.

The time to metal could be reduced if the foundry was included in early planning.12010080s r u o H6040200ActivityFigure 6.Figure 8. Mold Making Time EstimatesCost of composite moldsUsing this research as a baseline, the cost of producing composite tooling for patterns thatare 11”x11”x4” is fairly reasonable. Assuming that mold boxes and fixtures are reusable, and steam vents are both relatively inexpensive and functional, this process could result in prototype tooling for as little as $12,000. For this project, expenses reached almost $20,000, not including equipment development (mold-boxes/fixture/pressure-vessel/experiments). A large portion of these expenses were in RP pattern post-processing (repairs) and steam-vent machining, two tasks that can be easily refined.

A large portion of these expenses were in RP pattern post-processing (repairs) and steam-vent machining, two tasks that can be easily refined. With trained mold-makers and reduced machining (increased molded- in-place features) the cost in labor would bring total cost down substantially.ConclusionsThis research shows that molds can be produced within a two week time period withacceptable mechanical properties. Steam vents need to be improved and ejector pins should be512incorporated. Quality of patterns produced was not acceptable but very promising. The cost of the molds was reasonable, between $12,000-20,000, depending on geometry and size.

Quality of patterns produced was not acceptable but very promising. The cost of the molds was reasonable, between $12,000-20,000, depending on geometry and size. Benefits of using this type of rapid tooling include: • Cost of tooling is significantly less than production tooling • • • Early patterns can be used to design tree configurations • Early patterns and castings can be used to set-up and program production equipment • Early castings can be tested and redesigned several times before production begins • Composite tooling can potentially act as bridge tooling • Composite tooling can potentially become production tooling for short runsImproved time to market Improved quality of cast productsFuture RecommendationsAlthough patterns of castable quality did not result from this research, the feasibility of using this approach for polystyrene patterns was apparent. Several portions of this composite tooling process should be refined including: 1.

Several portions of this composite tooling process should be refined including: 1. Steam-Vent --design and production method 2. Use small scale molds to refine vent design as well as other process parameters 3. Refine basic fixture design for better mold attachment and part ejection 4. Mold box improvements for reduced machining (mounting-holes and datum-planes cast-in-place)5. Vacuum assist polystyrene bead fillAcknowledgementsWe gratefully acknowledge the support of Ford Motor Company and the RapidPrototyping Consortium Members.The authors would also like to thank the MSOE, Rapid Prototyping Center Students fortheir contribution on this project.References1. US Patent 4612968: Method of casting using expendable patterns 2. Hilton, Peter D., Paul Jacobs, “Rapid Tooling, Technologies and Industrial Applications,” Marcel Dekker, Inc, 20003. Atkinson, David, “Rapid Prototyping and Tooling,” Strategy Publications, 1997 4. Jacobs, Dr.

direct Selective Laser Sintered Fully Ferrous Components – Infiltration Modeling, Manufacturing and Evaluation of Mechanical Properties Phani Vallabhajosyula and David L. Bourell Laboratory for Freeform Fabrication, Advanced Manufacturing Center The University of Texas at Austin 1 University Station, MC C2200 Austin TX, 78712-0292 USA +512-471-3170 voice, +512-471-7681 fax phani.charana@gmail.com, dbourell@mail.utexas.eduAbstract Commercially available Selective Laser Sintered (SLS) ferrous components contain copper-based infiltrant in a ferrous preform. The choice of infiltrant has led to inferior mechanical properties of these components limiting their use in many non-injection-molding structural applications, particularly at elevated temperature. In the present work, an attempt has been made to replace the copper-based infiltrant considering cast iron as potential infiltrant for its fluidity, hardness and stability at comparatively high temperature.

A critical issue associated with the infiltration was diffusion of carbon from the cast iron into the steel preform thereby decreasing its melting temperature and distorting the part geometry. A predictive model was developed which defines the degree of success for infiltration based on final part geometry and depending on the relative density of the preform and infiltration temperature. The model may be extended to other ferrous powder and infiltrant compositions in an effort to optimize the properties and utility of the final infiltrated part. Initial studies were carried out to validate the model by infiltrating SLSed Laserformtm A6 tool steel preforms with ASTM Type I A532 cast iron. The parts were analyzed for geometry, microstructure and hardness. This research was sponsored by the National Science Foundation under Grant #DMI-0522176.1. Introduction:Selective Laser Sintering is a fast growing process for rapid prototyping and rapid manufacturing of functional parts.

Introduction:Selective Laser Sintering is a fast growing process for rapid prototyping and rapid manufacturing of functional parts. The process of selective laser sintering suffers from a common deficiency of not being able to produce pore-free parts from many materials. Since porosity diminishes material properties like hardness and ductility, it has to be eliminated as much as feasible. Metal infiltrated metal systems based on selective laser sintering (SLS) have been commercialized for some time [1-5]. Proper infiltrant selection can impart desirable properties to the final part in addition to preserving the part geometry [6, 7].! Selection of bronze infiltrant would lead to a compromise in strength and hardness of the part resulting from generally inferior mechanical properties of bronze compared to those of tool steel [8]. These parts can be used for applications where high thermal conductivity is desired such as short run injection moulds.

These parts can be used for applications where high thermal conductivity is desired such as short run injection moulds. High temperature stability of ferrous components demands a stable second phase in the ferrous matrix when compared to the commercial (copper based) infiltrant. New infiltrant metals have been considered for which hardness and maximum service temperature are not so severely compromised. In the present work, cast iron is considered as a potential infiltrant for its good fluidity and hardness. Cast iron is also available in a variety of strength grades and microstructures.This paper deals with modification to the commercial infiltration process whereby a low- melting-point cast iron is substituted for the copper alloy infiltrant. The methodology involves formation of green part as a result of selective laser sintering of LaserFormtm A6 tool steel powder [9-11], burnout of the binder and pre-densification of the component followed by infiltration with cast iron.

A critical issue is avoiding melting of the tool steel during infiltration. Diffusion of carbon from the cast iron into the tool steel preform results in lowering of the melting point of the tool steel part significantly. The work was carried out in three stages: constructing a predictive model depending on the relative density of the preform and infiltration temperature to define the degree of success for infiltration based on final part geometry, experimental part generation and analyzing the experimental results by correlating with the model constructed. The experimentation and the resulting parts validation for this model are also discussed in detail.2. Predictive model for infiltrationTo ensure the production of a decent infiltrated part, it is critical to set the infiltration temperature such that the solid fraction at equilibrium does not fall too much below the solid fraction of SLS part prior to infiltration.

Preferably the equilibrium solid fraction of the final infiltrated part will be greater than the uninfiltrated tool steel part relative density. The initial density of SLS part prior to infiltration can be varied by sintering the green part at various temperatures.a) Densification or Ashby plots:Once the green part is obtained by conventional SLS of the tool steel powder with binder, it is sintered in a vacuum furnace to drive off the binder and create a porous particle network [12]. During the initial stages of sintering coalescence of powder particles occurs by neck formation and growth between the contacting particles. During the intermediate stages particles tend to move closer to one another and the corresponding microstructure is characterized by a tunnel-network of porosity between the particles. Finally the network ceases to exist, giving rise to the formation of isolated spheroidal pores. These pores continue to shrink until the densification is complete.

Finally the network ceases to exist, giving rise to the formation of isolated spheroidal pores. These pores continue to shrink until the densification is complete. This ensures some hardness to the brown part. The resulting density of the component depends on the initial density, sintering temperature and dwell time at that temperature [13]. !the powder occurs by various mechanisms like plasticity, boundary diffusion, volume diffusion, Nabarro-Herring creep, Coble creep and Power law creep. [14, 15] defined Ashby, et al. densification for various relations mechanisms and created a Temperature Densification Map, a plot of relative density as a function of temperature T. The Ashby plot for densification of LaserFormtm A6 Tool steel with initial relative density of 0.48, approximately, is shown in Figure 1.

The Ashby plot for densification of LaserFormtm A6 Tool steel with initial relative density of 0.48, approximately, is shown in Figure 1. The resulting relative density for the parts sintered at 1100oC for 3hours and 1260oC for 5hrs is marked by points a, b respectively on the plot.Densification ofb!!a!!Fig.1. Ashby densification map for tool steel, plotting relative density as a function of temperature and time of sintering. The data points a,b represent increased relative densities of the parts sintered for 3hrs at 1100oC and 5hrs at 1260oC respectively.b) Equilibrium solid fraction at infiltration temperatureInfiltration into the brown part should ensure that cast iron melts completely and diffuses into the brown part without disturbing the part geometry. It was hypothesized that the equilibrium state at infiltration temperature is a mixture of solid tool steel and liquid cast iron.

It was hypothesized that the equilibrium state at infiltration temperature is a mixture of solid tool steel and liquid cast iron. A predictive model for infiltration of cast iron into selective laser sintered tool steel is constructed based on the fact that the equilibrium state at the infiltration temperature is governed by i) carbon content of tool steel and cast iron, ii) infiltration temperature, iii) relative density of tool steel SLS part and iv) amount of infiltrant (excess) needed to ensure complete infiltration.Let Ci and Cm be the carbon composition of cast respectively. tool steel Considering ! as the density of pure tool steel and cast iron, if " is the relative density of the brown tool steel part, the relative density of cast iron to be infiltrated is given by 1-".

Let Ce be the equilibrium carbon composition of the resulting part, which may be given by the relation,iron andCe= [Cm " +Ci (1- ")] = Ci- " (Ci-Cm) (1)As a reference, Figure 2 shows iron- carbon phase diagram with various materials compositions represented [16]. Let CL and CS be the carbon compositions of solidus and liquidus lines at equilibrium which in turn depend on infiltration temperature. The equilibrium solid fraction at the infiltration temperature is given by,Fig.2. Iron-carbon equilibrium diagram.

The equilibrium solid fraction at the infiltration temperature is given by,Fig.2. Iron-carbon equilibrium diagram. The definitions for carbon compositions at various points are given in the text.Xe=C CLL! !C CeS=[ CL!( C "+ i C-C LCiS!Cm] )Therefore, the equilibrium solid fraction (Xe) at equilibrium state is a function of relativedensity of the brown part (") and the infiltration temperature (Ti).This paper deals with infiltration of SLS tool steel part of carbon composition (Cm) 1.09% (wt) with ASTM A532 Class IA white cast iron of carbon composition (Ci) of 3.56% (wt). Based on Equation 2, a graph predicting the equilibrium solid fraction at various temperatures for different relative densities of the brown part is plotted as shown in Figure 3. The dashed line crossing the curves implies the equilibrium solid fraction is equal to the relative density of the brown part at a given infiltration temperature.

The dashed line crossing the curves implies the equilibrium solid fraction is equal to the relative density of the brown part at a given infiltration temperature. A point in Region I i.e., above the dashed line represents a higher value of solid fraction than the value prior to infiltration. The cast iron melts and infiltrates into the solid brown part resulting in a decent fully ferrous component. Any point below the dashed line i.e., in the Region II, represents the equilibrium solid fraction to be actually less than the relative density of the brown part. In this case, melting of the part is observed resulting in the distortion of the part.!(2)Region IRegion II! Fig.3. Equilibrium solid fraction as a function of infiltration temperature for different densities of brown part. The data points marked on the plot represent figures of various infiltrated parts; the details are discussed in the text.Fig.4.

The data points marked on the plot represent figures of various infiltrated parts; the details are discussed in the text.Fig.4. Effect of excess infiltrant on effective starting density for various densities of tool steel brown part. This also results in a decrease of infiltration equilibrium temperature as the amount of infiltrant increases.solidfractionatc) Effect of Excess InfiltrantExcess of infiltrant is needed to ensure complete infiltration of the part. Consider an excess of infiltrant (cast iron) be placed on the tabs of tool steel perform during infiltration. This actually alters the effective solid fraction, decreasing it initially. This in turn decreases the equilibrium solid fraction at infiltration temperature.

This actually alters the effective solid fraction, decreasing it initially. This in turn decreases the equilibrium solid fraction at infiltration temperature. Let # be the fraction of excess infiltrant by volume to the actual volume of infiltrant needed i.e., when # is equal to zero implies the absence of excess infiltrant, and when # equals to 0.2 implies the volume of infiltrant taken is 20% in excess to the actual volume of infiltrant needed. The effective starting density of green part as a function of excess infiltrant is given by,effÄ=Ä [ )Ä1)(å+1(+ÄÄ ])Ä-1(å-1= ] [Figure 4 shows the decrease in the effective starting density as # increases. The amount of decrease for a reasonable values of 0 < # < 0.5 results in only modest changes to "eff of less than 10% approximately.3.

The amount of decrease for a reasonable values of 0 < # < 0.5 results in only modest changes to "eff of less than 10% approximately.3. ExperimentalVarious experiments were carried out to validate the predictive model, in which cubes of dimensions 25mmx25mmx25mm, with a tab for placing the cast iron chunks were built on SLS machine from LaserFormtm A6 tool steel. The green density of these SLS parts was obtained from the ratio of weight of the part to its measured volume. Once the green density was known, the part was buried in alumina and put in a furnace to burn out the binder and pre-densify the part. The part was heated at 90oC/hr to a temperature of 750oC to burn out the binder and then at a rate of 100oC/hr to the set temperature, held at that temperature for certain time followed by furnace cooling. The theoretical relative density of the brown part could be predicted from the Ashby plot by knowing the set temperature and the dwell time.

The theoretical relative density of the brown part could be predicted from the Ashby plot by knowing the set temperature and the dwell time. This was compared with the actual relative density of the part calculated from its weight and volume.The pore volume was obtained from the relative density of brown part and the weight of cast iron needed for infiltration was determined. The tab on the brown part was carefully cleaned!(3)and cast iron chunks were placed on it for infiltration. The assembly was buried in alumina and placed in furnace. The heating rate was set at 100oC/hr up to the infiltration temperature and the sample was held there for 50 min followed by furnace cooling. Infiltration of the brown parts with different relative densities was carried at different temperatures to validate the above model.4. Testing the infiltrated SLS partsThe parts produced were machined to remove the tab portion and tested for any distortion in the part geometry.

Testing the infiltrated SLS partsThe parts produced were machined to remove the tab portion and tested for any distortion in the part geometry. Once the parts were found to be geometrically sound, they were then tested for geometry, density, microstructure, hardness, tensile strength and transverse rupture strength.a) Geometry and Density: The density of the part was determined using Archimedes method. For the infiltrated parts without any distortion, density was determined from the ratio of the part weight to its volume calculated from the dimensions. Once the density of the part was known, relative density can be found from the ratio of the actual density to the theoretical density of the fully dense part. (Theoretical density of the full dense part is ~ 7.8 g/cc [9])b) Microstructure: Microstructure of the infiltrated parts was obtained using optical microscopy.

(Theoretical density of the full dense part is ~ 7.8 g/cc [9])b) Microstructure: Microstructure of the infiltrated parts was obtained using optical microscopy. Microstructure also aided in determining the amount of porosity in the infiltrated parts.c) Hardness: Rockwell hardness C of the parts was determined. The surface of the parts was polished and indented with a Brale indentor. A preload of 10 kg was applied followed by a 150kg force. The deflection on the scale as indicated by the pointer gave the hardness of the infiltrated part. The hardness was measured at different points on the part to test for the consistency in the value. Knoops micro hardness was also found for different phases present in these parts using 500kgf load and equivalent Rockwell hardness was determined [17].5. Results and DiscussionVarious infiltrated parts were obtained from infiltrating brown parts with different initial relative densities at various infiltration temperatures.

Results and DiscussionVarious infiltrated parts were obtained from infiltrating brown parts with different initial relative densities at various infiltration temperatures. The geometry of the resulting parts is discussed.a) Geometry and densityGreen parts from SLS station with a relative density of 0.48 were pre-densified in vacuum at a temperature of 1260oC for 5hours, increasing the relative density to 0.62. Depending on this value of relative density, the weight of cast iron needed for infiltration was calculated. Both the part and the cast iron chunks placed on its tabs were buried in Alumina and set for infiltration at 1190oC. The part shown in Figure 5 is the decent infiltrated part obtained as a result. From the predictive model, it can be seen that at this infiltration temperature, the equilibrium solid fraction of the part is greater than 0.62 which implies no melting of brown part and thereby results in an acceptably infiltrated part.

The relative density of this part was found to be 0.8. Figure 6 shows another sample, with relative density of brown part being 0.57 and infiltrated at 1230oC. The resulting part with a relative density of 0.64 was sound, thus validating the predictive model.The part shown in Figure 7 represents a brown part with relative density of 0.5 infiltrated at 1275oC which resulted in melting of tool steel and distorted part geometry. It is evident that selection of infiltration temperature for the given brown part relative density from Region II of the predictive model (Figure 3) resulted in melting of tool steel and distorting the part. The part in Figure 8 also validates the model. Distortion in the final part is observed when a brown part with relative density of 0.52 was infiltrated with cast iron at a temperature of 1310oC.

The part in Figure 8 also validates the model. Distortion in the final part is observed when a brown part with relative density of 0.52 was infiltrated with cast iron at a temperature of 1310oC. Thepredictive model also explains the melting of tool steel as the equilibrium solid fraction at this temperature is found to be less than 0.52.!!!!! !!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!!Fig.5. SLSed A6 tool steel part pre-sintered from a relative density of 0.48 to 0.62 by heating to 1260oC for 5hrs and then infiltrated at 1190oC. The resultant density of infiltrated part was 0.8!Fig.6. SLSed A6 tool steel part pre-sintered from a relative density of 0.48 to 0.57 by heating to 1260oC for 5hrs and then infiltrated at 1230oC. The resultant density of infiltrated part was 0.64!!Fig.7. SLSed A6 tool steel part pre-sintered from a relative density of 0.49 to 0.5 by heating to 1160oC for 3hrs and then infiltrated at 1275oC.

SLSed A6 tool steel part pre-sintered from a relative density of 0.49 to 0.5 by heating to 1160oC for 3hrs and then infiltrated at 1275oC. Melting of tool stell and hence distortion of the specimen is observed as the infiltration temperature for given brown density lies in Region II!Fig.8. SLSed A6 tool steel part pre-sintered from a relative density of 0.46 to 0.52 by heating to 1100oC for 3hrs and then infiltrated at 1310oC. As the infiltration temperature for given brown density lies in Region II, melting of tool steel along with cast iron is observed, distorting the final part. !b) MicrostructureTo get an insight into the effect of infiltration on laser sintered parts and the amount of porosity present in these parts, microstructural analysis was carried out. The parts were fine polished and etched with 2% nital for 20 seconds to obtain the microstructure.Fig.9. Presence of porosity can be observed from the micrograph (5X) of an unetched component with " = 0.8.Fig.10.

Presence of porosity can be observed from the micrograph (5X) of an unetched component with " = 0.8.Fig.10. The microstructure of an infiltrated component with "=0.8 indicating of the porosity and martensite. (10X)presenceFig.11. The presence of acicular martensite the evident microstructure of an infiltrated component with "=0.8 (20X)!fromConsiderable porosity is observed from the micrograph of component with relative density 0.8 as shown in Figure 9 and Figure 10. The optical micrographs shown in Figure 10 and Figure 11 indicate acicular martensite in the microstructure. Fine network of cementite is observed to surround the prior formed austenite grains.

The optical micrographs shown in Figure 10 and Figure 11 indicate acicular martensite in the microstructure. Fine network of cementite is observed to surround the prior formed austenite grains. The presence of martensite imparts hardness to the component, while porosity decreases it significantly.c) HardnessRockwell hardness of the polished parts is obtained, and from the hardness values it can be inferred that hardness of the component seems to depend on its relative density, in other words, the amount of porosity. By approaching higher densities for the infiltrated component, it can be possible to achieve higher hardness comparable with that of ASTM A532 white cast iron (53HRC, ASTM Standard).The general dependence of strength on porosity is given by $ = K $o (1- #)m = K $o (")m where $ is the strength, K and m are constants [18, 19]. The relationship can generally be applied to yield strength, tensile strength and three or four point bend strength.

The relationship can generally be applied to yield strength, tensile strength and three or four point bend strength. The plot shown in Figure 11 is obtained from Equation 4 and using the porosity and corresponding BHN linear values observed for relationship implies that the hardness also varies exponentially with relative density with an exponent of 1.5.!(4)the parts. TheH = 416.87 (")1.5Fig.11. Plot for hardness (BHN) values of infiltrated parts versus linear relationship implies that hardness varied linearly with relative density with an exponent of 1.5.!their relative density. TheMicro-hardness: Due to the presence of high porosity, the infiltrated components were tested for micro hardness. Knoops hardness was measured using load of 500kgf. In the region with high density of acicular martensite, the hardness was observed to be 616KHN which is equivalent to 54HRC and the hardness at other region was found to range from 436-500KHN , equivalent to 43-47.2HRC.

The high hardness value in the later region can be attributed to the presence of bainite in the microstructure.6. ConclusionsA predictive model for infiltration of cast iron into SLS tool steel part was developed and tested for different relative densities of brown parts and at various infiltration temperatures. From the above experiments, it was shown that final parts were consistent with the predictive model, and that fully ferrous components can be obtained by proper selection of infiltration temperature for a given relative density of brown part.The parts obtained should be tested for other mechanical properties like strength and ductility. The predictive model can be further tested considering infiltration with excess infiltrant. Fully dense components might be produced by increasing the number of tabs attached to the part or by adding excess amount of cast iron.

Fully dense components might be produced by increasing the number of tabs attached to the part or by adding excess amount of cast iron. The predictive model can be developed further for different ferrous alloys and validated by repeating the experiments at different sintering and infiltration temperatures. The methodology may be tested by first choosing theproduct composition for a featured application and working back to determine the required matrix and infiltrant compositions.ACKNOWLEDGEMENTSThis research was sponsored by the National Science Foundation Grant DMI-0522176REFERENCES1. J.J. Beaman, J.W. Barlow, D.L. Bourell, R.H. Crawford,H.L. Marcus, K.P. McAlea, "Solid freeform Fabrication: A New Direction In Manufacturing", Kluwer Academic Press, Boston, 1997.2. R.B. Heady, J.W. Cahn, “An Analysis of the Capillary Forces in Liquid-Phase Sintering Of Spherical Particles”, Met Trans, 1970, 1#1, 185-189.3. F.H.

DIRECT TISSUE SCAFFOLD FABRICATION VIA FUSED DEPOSITION MODELING AND BIOMIMETIC MINERALIZATIONJesse Bernardo1, Satyavrata Samavedi2, Christopher B. Williams1,3, Abby R. Whittington2,4 1 Department of Mechanical Engineering, 2 Department of Chemical Engineering, 3 Department of Engineering Education, 4 Department of Materials Science and Engineering Virginia Polytechnic Institute and State UniversityABSTRACTTo alleviate material limitations of the additive manufacture of tissue scaffolds, researchers have looked to indirect fabrication approaches. The feature resolution of these processes is limited however, due to the viscous ceramic slurries that are typically employed. To alleviate these limitations, the authors look to an indirect fabrication process wherein a pattern, created using Fused Deposition Modeling, is biomimetically mineralized with an aqueous simulated body fluid, which forms a bonelike hydroxyapatite throughout the scaffold pattern.

Mineralized patters are then heat treated to pyrolyze the pattern and sinter the minerals. With this process, scaffolds were created with wall thicknesses as small as 150 m and internal channel diameters of 280-340 m, an appropriate range for bone tissue engineering.Keywords: Fused Deposition Modeling, Biomimetic Mineralization, Tissue Scaffold, Additive Manufacturing1. TISSUE SCAFFOLDS WITH DESIGNED MESOSTRUCTUREA tissue scaffold is an artificial cellular structure composed of a biocompatible material that mimics a body’s extra-cellular matrix (ECM) and supports three-dimensional tissue formation in order to begin healing an injury [1- 3]. The scaffold’s internal structure (featuring interconnected pores and channels) provides surface area for cell attachment and pathways for cell movement [4, 5], surface area for cell attachment, diffusion of nutrients, and removal of cell wastes within the scaffold.

In addition, it provides mechanical and chemical cues that affect the cellular ingrowth, differentiation, and vascularization [6-8].The mesostructure of the scaffold (i.e., the size, orientation and geometry of the scaffold pores) can affect its ability to mimic the ECM and thus affects how well cells proliferate throughout the structure [9-11]. Several researchers have found that different cell types prefer different scaffold structures, with porosity and pore size being the most commonly benchmarked properties. Scaffolds with a high porosity offer more “living space” for cells to move into and start to grow. Scaffold porosity should be greater than 75% to ensure cell proliferation [9]; however, the maximum porosity is limited by the required mechanical strength of the scaffold – a key tradeoff in the design and fabrication of tissue scaffolds.

Furthermore, pores that are too small will become clogged and cause a decrease in nutrient diffusion, cellular penetration, ECM production, neovascularization [4] and possibly cell death; pores that are too large form gaps that the cells cannot cross. Pore sizes on the range of 200-400 micrometers are accepted for bone cell proliferation [1, 12, 13].Although porosity and pore size provide the living space and pathways for cell movement, the scaffold’s permeability affects how nutrients are delivered to, and wastes are removed from, the cells. Ku et al. suggest that permeability is a combination of several scaffold factors including: porosity, pore size, distribution and orientation, and interconnectivity [14]. The tissue scaffolds’ permeability should provide nutrients until a vascular system is in place [11].

The tissue scaffolds’ permeability should provide nutrients until a vascular system is in place [11]. Once a vascular system exists throughout the scaffold, cell nutrient and waste removal is carried through the blood vessels.Microscale topology can also affect cell behavior within the scaffold. Specifically, surface topography will influence cellular attachment, movement and cellular signaling [15].

Specifically, surface topography will influence cellular attachment, movement and cellular signaling [15]. The scaffold’s bulk crystallinity [16], surface roughness [4] and microporosity [10] are known to affect cellular attachment, proliferation and migration and neovascularization. Corresponding Author: 114F Randolph Hall, Blacksburg, VA, 24061; cbwilliams@vt.edu; (540) 231-34221Finally, the mechanical properties of the scaffold should match those of the tissue it is replacing; this will protect the cells from mechanical loads, contractile forces from tissue formation [17], and any in vivo/vitro stresses [15, 18].Thus, scaffold structure should be designed to support specific cell types in order to maximize the growth potential of the desired tissue [8, 15]. However, existing scaffold fabrication techniques do not permit this level of design freedom. While stochastic fabrication techniques (e.g.

However, existing scaffold fabrication techniques do not permit this level of design freedom. While stochastic fabrication techniques (e.g. solvent casting and particle leaching [19], gas foaming [20, 21], and emulsification, freeze-drying and thermally induced phase separation [22-24]) processes result in highly porous scaffolds, the size, shape, location and orientation of the porosity is stochastic, and thus is unable to provide repeatable results. Furthermore, these methods are process dependent and do not allow for the strategic placement of material, and therefore do not allow for control over pore placement and internal architecture.Additive Manufacturing (AM) technologies’ precise control of pore size, geometry, interconnectivity and distribution throughout the scaffold affords the creation of scaffold mesostructures that are designed for specific tissues [25].

Many different AM methods have been used to create biocompatible scaffolds (as reviewed in the authors’ prior work [26]). 3D Printing has been used to create PLGA scaffolds [27, 28]. Scaffolds made of ceramics (e.g., hydroxyapatite [29] and calcium phosphate [30]), biodegradable polymers [31], and hydrogels [32] have been created by stereolithography (SL) techniques. Laser sintering (LS), which uses a laser to selectively fuse together powder particles in a powder bed, has been used to create scaffolds made of ceramic/polymer blends such as polymer-coated calcium phosphate [33, 34] and hydroxyapatite [35].While each of these AM technologies is capable of producing scaffolds with controlled geometry and porosity with repeatable results, they do suffer from some processing limitations.

The primary limitation is the relatively few biocompatible materials that are compatible with each system – materials available to AM techniques are mostly constrained to polymers that are specialized for each process. For example, SL scaffold fabrication is limited to materials that can be cured by UV irradiation, thus limiting its ability to directly fabricate scaffolds for hard tissue [36]. Extrusion processes are limited to materials that can be melted into a semi-liquid phase before extrusion [37]. Many researchers have tried to extend the processes’ working materials by integrating bio-compatible ceramic particles into polymer systems as in LS [38, 39], SL [40-42], and Robocasting [43, 44].

Many researchers have tried to extend the processes’ working materials by integrating bio-compatible ceramic particles into polymer systems as in LS [38, 39], SL [40-42], and Robocasting [43, 44]. However, the resultant parts suffer from a reduced resolution and pore size [41, 45], and delamination, cracking and warping as the large concentrations of polymeric binder are pyrolyzed [40, 46, 47].To bypass these material limitations, indirect fabrication techniques (i.e., investment casting approaches) have been proposed; however, their reliance on viscous ceramic slurries prevents the realization of scaffold geometries with fine features (Section 2). In this paper the authors detail their exploration of an indirect scaffold fabrication process wherein biomimetic ceramic mineralization is used to coat scaffold patterns that are fabricated by additive manufacturing. The process alleviates both the material and geometric limitations imposed by existing scaffold fabrication techniques.

The process alleviates both the material and geometric limitations imposed by existing scaffold fabrication techniques. An overview of the indirect fabrication process is provided in Section 3. Experimental methods are described in Section 4. In Section 5, the results of process experimentation are provided. Finally, closure and opportunities for future work are described in Section 6.2. INDIRECT FABRICATION OF SCAFFOLDSIndirect scaffold fabrication techniques have been explored as a means of circumventing the inherent material processing limitations of existing AM techniques. Existing indirect scaffold fabrication techniques typically follow a strategy similar to investment casting. First, a negative mold of the desired scaffold is fabricated, into which a biocompatible slurry is cast. The slurry contains the scaffold material (e.g., HA, PLGA, collagen), solvents (e.g., chloroform and methanol), and porogens [48-50]. The slurry is then solidified via a curing process.

The slurry contains the scaffold material (e.g., HA, PLGA, collagen), solvents (e.g., chloroform and methanol), and porogens [48-50]. The slurry is then solidified via a curing process. After the slurry has set, the mold is removed – leaving behind a biocompatible scaffold with internal channels.Hydroxyapatite (HA) macro-porous scaffolds have been created by using SLA [49] and 3DP [51] to create negative molds. After mold fabrication, a HA slurry is cast and dried in the mold. The cast is heated to remove mold material and then sintered to increase the density of the HA. The wall thicknesses created from the mold process were not smaller than other AM techniques: SLA molds were ; 3DP molds were . In addition, these HA scaffolds did not contain a “local” pore distribution.Sachlos et al. fabricated porous biocompatible collagen scaffolds with channels by casting a collagen/water solution into a mold printed by a Model Maker II (Figure 1) [52].

fabricated porous biocompatible collagen scaffolds with channels by casting a collagen/water solution into a mold printed by a Model Maker II (Figure 1) [52]. No new material formulation for printing the mold was necessary; the machine manufacturer’s printing mediums (ProtoBuildTM and ProtoSupportTM) were used. After fabricating the mold, the ProtoSupportTM material is removed by submerging the mold in a solvent (BioActTM) to create a negative mold. The collagen suspension was cast into the mold and frozen to form a tissue2scaffold. The mold and scaffold were then immersed in ethanol, which dissolved the mold material and ice crystals. Critical point drying was used to remove any remaining solvents and ice crystals. Shrinkage during critical point drying reduced the accuracy of the printed channels.

Critical point drying was used to remove any remaining solvents and ice crystals. Shrinkage during critical point drying reduced the accuracy of the printed channels. An increase in collagen dispersion in the cast suspension decreases shrinkage [52] but is limited because an increase in viscosity can prevent the collagen suspension from penetrating small channels. Although some scaffolds were designed to have a channel width of , the smallest channel widths reported for this processes were .Figure 1. A) Designed channel width. B) Printed mold with varying channel widths. C) Porous collagen mold with designed channel widths. [52].In a similar manner, Yeong et al. fabricated biocompatible collagen scaffolds by casting a collagen/water solution into a mold printed by a inkjet printer (Solidscape’s T612 Benchtop) [50]. No new formulation of mold material was needed and the machine’s original printing mediums (InduraCast and Indurafill) were used to print the mold.

No new formulation of mold material was needed and the machine’s original printing mediums (InduraCast and Indurafill) were used to print the mold. The collagen suspension was frozen in the mold and then immersed in ethanol to dissolve the mold [53]. The scaffold was then freeze-dried to remove water crystals from the collagen suspension leaving behind randomly oriented pores [53]. A branched channel design with channel widths of 200 and 300 was achieved.Lee et al. fabricated collagen scaffolds with “villa” like features [54] using 3DP to create a plaster mold. The plaster mold was infiltrated with PEG to increase its strength before the mixture of PLGA, chloroform, methanol and sucrose was cast. The PLGA cast was dried in a fume hood before being immersed in deionized water. Water dissolved the mold and sucrose (which acted as a porogen) to leave behind porous “villa” like features that tapered from in diameter and high.

Water dissolved the mold and sucrose (which acted as a porogen) to leave behind porous “villa” like features that tapered from in diameter and high. The pores that formed from leaching sucrose were uniformly distributed and well connected [54].Taboas et al. [55] fabricated “multi-pore-architecture” poly(L)lactide (PLA) scaffolds that contained controlled macro-pores (interconnected macro-pores in diameter) with local porosity created by particle leaching (Figure 2). Macro-pores may increase diffusion of nutrients to cells deep in the scaffold while local pores provide living space for the cells [55].Figure 2. A) PLA scaffold with designed macro-pores and local pores. B) Designed macro-pores are interconnected (in a three-dimensional structure) cylindrical channels.

A) PLA scaffold with designed macro-pores and local pores. B) Designed macro-pores are interconnected (in a three-dimensional structure) cylindrical channels. C) SEM of local porosity created by salt leaching [55].The use of AM to manufacture molds presents three important contributions to the design of tissue scaffolds:  Using molds relaxes the material limitations associated with AM. The mold material does not have to be biocompatible, thus scaffolds can be fabricated “from the largest pool of materials available [55].” Also, materials that are already compatible for the AM machine can be used instead of developing new materials [51].Allows for the simultaneous creation of macro-pores (channels) and local pores (porogen) [55].

AM molds can produce scaffolds with designed internal channels and may increase cell viability in solvent-cast/porogen-leach scaffolds [52] by increasing the nutrient flow to cells in the scaffold.Toxic/residual materials can be removed more easily from the scaffold. Whereas “solvent casting and particle leaching” methods were limited by thickness, indirect methods contain macro-pores which enhance the removal3of mold material. Macro-pores allow the removal of toxic solvents, and porogens from the scaffold allowing scaffolds to be fabricated with thicker cross sections [55].Typical restrictions of traditional casting processes, imposed by an inability to fabricate molds that would result in complex geometry and internal architecture, are alleviated by fabricating molds via AM [56]. However, existing indirect fabrication approaches do limit the features that are able to be produced. In general, the processes are limited by:  Non-cell friendly separation of mold from the cast.

However, existing indirect fabrication approaches do limit the features that are able to be produced. In general, the processes are limited by:  Non-cell friendly separation of mold from the cast. The mold materials must be completely removed from the cast material without compromising the scaffold architecture. Harsh chemicals and extreme temperatures can remove the mold while leaving the cast intact but prohibit the incorporation of living cells or growth hormones.Weak mechanical properties due to the creation of local pores within the scaffold architecture.  Mold imperfections are transferred to the cast [55].  Increased fabrication time. Scaffolds take longer to make than direct-fabrication AM because a mold must be fabricated, cast, and dissolved.Feature resolution is limited by the viscosity of the cast material. Viscous materials, such as high loaded ceramic suspensions, may trap air bubbles within the mold and cause defects in the scaffold architecture [53].

Viscous materials, such as high loaded ceramic suspensions, may trap air bubbles within the mold and cause defects in the scaffold architecture [53]. As a result of the viscosity limitations the internal channels and scaffold walls are larger than other AM produced scaffolds.Alleviating the existing limitations of indirect scaffold fabrication approaches is of merit, as doing so will allow the realization of designed scaffold structures that are not constrained to specific geometries or materials.3. INDIRECT FABRICATION OF SCAFFOLDS VIA ADDITIVE MANUFACTURING AND BIOMIMETIC MINERALIZATIONWhile indirect AM approaches to scaffold fabrication alleviate limitations imposed by traditional stochastic fabrication techniques (e.g., non-repeatability) and direct AM techniques (e.g., material selection), their use of viscous ceramic slurries constrains the geometries that are able to be realized.

To mitigate scaffold feature-size limitations, we look to biomimetic mineralization as a means of coating polymeric scaffold patterns with a bioceramic.Applying strategies found in biomineralization – the process by which living organisms secrete inorganic minerals to harden or stiffen existing tissues in the forms of skeletons, shells, teeth, etc. – biomimetic mineralization is the formation of new crystalline phase via nucleation followed by crystal growth in aqueous solutions at ambient conditions [57]. Biomimetic approaches to mineralization typically involve soaking the target in simulated body fluid (SBF) – an aqueous solution of salts that contains ions approximating those of human plasma [58]. SBF is influencing the mineral precipitates [59]. supersaturated with respect to apatite, with the concentration of SBF has been used to deposit formations of bonelike apatite, which is essential for material to bond to living bone [60].

supersaturated with respect to apatite, with the concentration of SBF has been used to deposit formations of bonelike apatite, which is essential for material to bond to living bone [60]. In addition, SBF coatings have good resorption [61] and have been shown to affect osteoblast viability, proliferation, and gene expression [62]. Coated implants and scaffolds have increased bone ingrowth (compared to non-treated implant) [63], biocompatibility [64], osteoconductive properties [58] and bone bonding [65]. In addition to creating a bone like apatite, and creating favorable conditions for bone growth, SBF has been used to mineralize several different materials: PLLA and PLGA [66], collagen [67], Ti6A14V [68], Bioglass® [69].The process chain is as follows (Figure 3):1. Fabrication of scaffold pattern: A scaffold pattern with designed pore geometries is fabricated using additive manufacturing.2.

Fabrication of scaffold pattern: A scaffold pattern with designed pore geometries is fabricated using additive manufacturing.2. Surface treatment of scaffold patterns: Scaffold patterns are chemically treated to activate their surfaces to increase nucleation sites and to enhance the rate and quality of the mineral deposition.3. Mineralization of scaffold patterns: The scaffold pattern is submerged in a simulated body fluid solution that deposits layers of biocompatible minerals (hydroxyapatite, HA) on its surface.4. Burnout and sintering of scaffold patterns: The mineralized scaffold pattern is pyrolyzed to remove the pattern material and to sinter the deposited minerals, thus leaving behind a three-dimensional bioceramic scaffold.4(b) Figure 3.

Indirect Scaffold Fabrication Process; a) scaffold pattern is fabricated by AM, b) scaffold pattern coated with bioceramic following mineralization, c) biocompatible scaffold formed after pyrolyzing pattern material(a)(c)Like existing indirect AM scaffold fabrication techniques, this approach will alleviate geometric limitations typical of stochastic methods and the material limitations of direct AM methods. The biomimetic mineralization coating approach, however, uses a low-viscosity aqueous SBF solution, which the authors hypothesize will better penetrate complex scaffold geometries and thus eliminate the geometry and feature-size limitations found in existing indirect techniques.4.

EXPERIMENTAL METHODS4.1 Pattern Fabrication via Fused Deposition ModelingExtrusion-based processes, such as Fused Deposition Modeling (FDM), are popular choices for AM scaffold fabrication as the biopolymers that they selectively extrude are sufficiently stiff to span small gaps between scaffold struts, and thus eliminate the need for a secondary support material [25]. Scaffold porosity topology is altered by varying the thickness of the deposited lines (road width, RW), the offset distance between lines (gap width, GW), and the offset angle between each layer (Figure 3a).In this work, the authors investigate two materials for the creation of scaffold materials: ABS (Stratasys P-400) and investment casting wax (ICW) material (Stratasys ICW-05).

These are not necessarily ideal pattern materials; an ideal material would be receptive for the mineralization process and would be suitable for removal via pyrolysis (i.e., is able to be removed via a heat treatment without significant warping and toxic ash residue). However, they are sufficient for providing a proof of concept for the process. In addition, these two materials were shown to span small gaps and have been used successfully in other burnout processes [70-72].Scaffold patterns are fabricated using a Stratasys FDM 1500 (ABS) and a Stratasys FDM 1600 (ICW) and a T12 nozzle (0.305 mm diameter). The scaffold fabrication printing parameters (i.e., road width (RW), road orientation, temperatures, etc.) used in experimentation are summarized in Table 1. Printed scaffold patterns were 12 mm in diameter, 1.5 mm in height and printed with a layer thickness of 0.30 mm (5 total layers).Table 1.

Printed scaffold patterns were 12 mm in diameter, 1.5 mm in height and printed with a layer thickness of 0.30 mm (5 total layers).Table 1. Pattern fabrication processing parametersMaterialABS ICWProcessing Temperature (Model/Support/Envelope) 270 / 265 / 270 °C 70 / 71 / 28 °CRoad Width (mm)0.3 0.4Road Orientation0-90° 0-90°Printing PatternSingle roads Continuous4.2 Scaffold Surface TreatmentMineralization via SBF does not typically result in cohesive coatings without a preliminary surface treatment of the polymeric substrate [26]. As such, the pattern must first undergo a surface treatment prior to mineralization. The authors investigated two candidate surface treatments:  Plasma treatment activates surface groups for mineralization and decreases the contact angle allowing for aqueous solutions to easily penetrate the scaffold [73]. Plasma treatment has been successfully used in stochastic scaffold processes including PCL [74], and PCL-HA [73].

Plasma treatment has been successfully used in stochastic scaffold processes including PCL [74], and PCL-HA [73]. Samples are plasma treated in an oxygen environment for 15 minutes at 50 milliamperes in an SPI Plasma Prep II.5Sodium hydroxide (NaOH) etching creates negatively charged groups on surfaces, and has shown to increase mineralization of treated surfaces when compared to identical, non-treated surfaces [58]. In this procedure, samples are immersed in 0.1M NaOH solution for 15 minutes.4.3 Scaffold MineralizationAs described in Section 3, simulated body fluid (SBF) is used to biomimetically mineralize the scaffold patterns. A 10x concentrated SBF formulation (Table 2) is used in this process since higher concentrations accelerate the formation of minerals and are stable for long periods of time [59, 75].

A 10x concentrated SBF formulation (Table 2) is used in this process since higher concentrations accelerate the formation of minerals and are stable for long periods of time [59, 75]. 1x and 5x concentrations were explored in preliminary experiments, but did not result in sufficient mineral coating on the patterns’ surfaces [26].To prevent premature precipitation, two aqueous solutions are prepared: a “stock” solution containing all of the ions except for the bicarbonate, and another aqueous solution containing only the bicarbonate (Table 3). Precipitation occurs when the two solutions are mixed (10x SBF is prepared from the stock and bicarbonate solution in a 4:1 ratio). Scaffold patterns are then anchored to the center of a petri-dish containing mixed SBF. Fresh SBF is provided every two hours to ensure that ion concentration remains at a level that produces mineralization.Table 2. Ion Concentrations (mM) of blood plasma and 10xSBFTable 3.

Fresh SBF is provided every two hours to ensure that ion concentration remains at a level that produces mineralization.Table 2. Ion Concentrations (mM) of blood plasma and 10xSBFTable 3. Salt masses for stock and bicarbonatesolutions.IonBlood Plasma 142 5.0 103 2.5 1.05 1.0 0.5 27.0 7.2-7.410xSBF [74] 1030.0 5.0 1065 25 5 10.0 -- 10.0 6.1Stock SolutionBicarbonate SolutionSaltsKClGrams (cid:10) 1.23 3.73 3.68 3.05 1.20 0.710.84pHTwo flow conditions (dynamic and static) are investigated to identify which would result in increased mineral deposits. Dynamic flows are tested to help mineralize the porous scaffolds by forcing fluid through them. For dynamic flow tests, the samples are placed on an orbital shaker (~70 Hz) inside an incubation chamber set at 37 °C.4.4 Pattern Pyrolysis and SinteringOnce the scaffolds are successfully mineralized, they are placed in a furnace to burn away the polymeric pattern material (ABS or ICW).

The heating profile used for ABS patterns is based on thermogravimetric analysis (TGA) of ABS by ToolworX and Stratasys Inc. [76]. Through this analysis it was found that ~95% of ABS is removed at ~450 °C, while the remaining material removed at 575 °C [71]. Similarly, the heating profile for Stratasys ICW-06 was based on research on bioceramic implants that used ICW-06 as a binder in a FDC processes. Bose et al. heated bioceramic implants up to 550 °C to evaporate the mold and binder and chose a slow heating rate of 2°C/min to avoid cracking or distortion of the part [77]. A slow heating rate (0.5 °C /min) was used to prevent cracking and distortion of the part.The sintering temperature chosen for both ABS and ICW samples in this study is 800 °C (with a 1 hour hold) ata rate of 0.5 °C/min.5.

RESULTS AND DISCUSSION5.1 Determining Surface Treatment and Flow Conditions for Scaffold MineralizationCritical to the development of this indirect scaffold fabrication process, is the identification of mineralization processing parameters that significantly increase mineral deposition. With SBF concentration determined, the surface treatment and flow type during mineralization were treated as variables to identify the combination which provided the best mineralization. A total of 8 scaffold samples, featuring permutations of ABS and ICW patterns, Plasma and NaOH surface treatments, and dynamic and static flows, were created and weighed before and after mineralization to determine the amount of materials deposited. The final mass of deposited minerals for each scaffold is presented in Table 4.6A three-way ANOVA was performed on the mineral mass data [78].

The final mass of deposited minerals for each scaffold is presented in Table 4.6A three-way ANOVA was performed on the mineral mass data [78]. The analysis indicated that flow condition and surface treatment were the two most significant factors in this experiment (F-ratios of 170 and 10.2), while the scaffold material had only a small effect (F-ratio of 1.89). The calculated p-values, which indicate confidence in our hypothesis, were 0.0486 and 0.194 (for the flow condition and surface treatment parameters, respectively), and indicate a strong relation between the experimental factors and amount of mineral deposition.Table 4. Mineral mass data (in grams)ABSICWDynamic0.01530.01080.01410.0109Static0.00330.00300.00260.0004As can be seen in Table 4, scaffolds that were plasma treated mineralized more than those treated with NaOH. Also, dynamic conditions resulted in much better mineralization than static conditions (a 77.2-96.3% increase).

Also, dynamic conditions resulted in much better mineralization than static conditions (a 77.2-96.3% increase). As such, all further experiments were conducted using the plasma surface treatment and dynamic flow conditions.5.2 Process ResultsTo validate the presented approach, patterns were fabricated from ABS and ICW with gap widths (GW) of 500 m. Following mineralization and sintering, these GW measurements will correspond to the wall thickness (WT) of the scaffold (Figure 3). Similarly, the internal channel diameter (CD) will be formed when the RW material is removed during sintering and will have a diameter of ~300 m and ~400 m for ABS and ICW patterns respectively (Table 1). Sample printed patterns are presented in Figure 4. As can be seen, the diameter of the RW for ICW is not constant.

Sample printed patterns are presented in Figure 4. As can be seen, the diameter of the RW for ICW is not constant. The low viscosity of the extruded wax causes the roads to sag when spanning gaps; roads swell (when supported) and thin (when unsupported), creating a wave-like pattern along the length of their deposition.Figure 4. A) ABS and B) ICW scaffold patterns with a 500 m gap widthABS and ICW scaffolds were plasma treated (Section 4.2) and then mineralized in a 10xSBF solution for 48 hours under dynamic flow conditions. Fracture surfaces of mineralized ABS and ICW samples are presented in Figure 5.Figure 5.

Fracture surfaces of mineralized ABS and ICW samples are presented in Figure 5.Figure 5. Fracture surface of pre-sintered scaffolds with GW = 500 m; (a) ABS, (b) ICWFollowing mineralization, scaffolds were dried overnight and then sintered to 800 °C at a rate of 0.5 °C/min.The mass of the samples were recorded before and after mineralization, and after sintering, to determine the amount of mineral deposition and to verify that the polymer pattern had been removed. Fracture surfaces of sintered samples with a GW = 500 m are presented in Figure 6. As can be seen, the patterns’ original pores (now the scaffold walls) are filled with mineral as desired.7PlasmaNaOHPlasmaNaOHFigure 6. Fracture surface of sintered scaffolds with GW = 500 m; (a) ABS, (b) ICWMeasurements of wall thickness (WT) and channel diameter (CD) were made following printing and following mineralization and sintering. Measurements of the samples with 500 m are presented in Table 5.

Measurements of the samples with 500 m are presented in Table 5. The “Pre- Sintered” values fall within the range of expected values (i.e., the “Pre-Mineralized” values), except for the ICW WT. Plasma treatment may have caused this area of the scaffold to deform, resulting in a smaller reported WT. A slight shrinkage during the burnout and sintering process is noted, which causes the CD to decrease slightly.Table 5. Wall thickness and channel diameter measurements for samples with GW = 500 mDesignedPrintedMineralized and SinteredGW / WT (m) RW / CD (m) GW / WT (m) RW / CD (m) WT (m)CD (m)ABS ICW500 500300 400510 430310 360530 470280 340To explore the hypothesis that mineralization via an aqueous SBF solution would be able to create finer scaffold features than existing indirect processes, additional ICW samples with gap widths of 150 m and 300 m were fabricated.

As can be seen in the images from field-emission scanning electron microscopy (FE-SEM) of the fracture surface of the scaffold, the SBF mineralization approach is able to deposit minerals in gaps as small as 150 m (Figure 7). The resulting wall thickness from these scaffolds was 158 ± 33.4 m.Figure 7. FE-SEM images of mineralized scaffold with wall thickness = 150 mTo benchmark the presented method against existing indirect fabrication methods, Detsch and coauthors’ scaffold fabrication process (highlighted row of Table 6) was selected as a benchmark since it has produced the smallest WT and CT features thus far in the literature. All fabrication methods were compared to the benchmark CD and WT values and assigned either a “0,” “+,” or “–” based on if they had an equal, better or worse feature resolution (Table 6).

All fabrication methods were compared to the benchmark CD and WT values and assigned either a “0,” “+,” or “–” based on if they had an equal, better or worse feature resolution (Table 6). The WT and CD values obtained from the fabrication process presented here (first two rows of Table 6) met or exceeded the resolution of other indirect methods.8Table 6.

The WT and CD values obtained from the fabrication process presented here (first two rows of Table 6) met or exceeded the resolution of other indirect methods.8Table 6. Comparison of WT and CD across existing indirect scaffold fabrication processesMold Fabrication FDM: FDM1500Mold Material ABSCast Material 10xSBFMeasured† CD WT 280 530Compared WT –CD +ReferenceFigure 6FDM: FDM1600ICW10xSBF158340+–Figure 7D-IJP: Solidscape MM2 SLA: 3D Systems D-IJP: Solidscape BT66WaxCibatool SL5170WaxPGA, PLA, HAHA slurryHA slurry~500560- 755~500500- 800 366- 968~3000–0––0[55][49][51]† All measurements are in micrometersTo chemically characterize the fabricated scaffolds, energy-dispersive X-ray spectroscopy (EDS) was completed on 12 different spots on the 150 m samples. EDS detected elements in all samples of calcium and phosphorous, with an average Ca/P ratio = 1.7 + 0.52.

EDS detected elements in all samples of calcium and phosphorous, with an average Ca/P ratio = 1.7 + 0.52. While it is difficult to speculate on the nature of mineral based on Ca/P ratio alone, it can be concluded with certainty that the mineral consists of calcium and phosphorous only, as other elements like magnesium and sodium were not detected. A Ca/P ratio of 1.7 indicates that it is slightly higher than that for hydroxyapatite (1.67), but is quite comparable. Therefore, we can speculate that the mineral formed is a type of hydroxyapatite, is probably crystalline in nature, and is not very soluble in aqueous environments. This also suggests that the surfaces may be osteoconductive and may promote growth and proliferation of bone-like cells. However, X-ray diffraction is needed to confirm the results of EDS.6.

This also suggests that the surfaces may be osteoconductive and may promote growth and proliferation of bone-like cells. However, X-ray diffraction is needed to confirm the results of EDS.6. CLOSURE AND FUTURE WORKIn this work, the authors present an indirect scaffold fabrication process similar to investment casting wherein polymer scaffold patterns are coated in a bioceramic via a biomimetic mineralization process. The patterns are first fabricated by Fused Deposition Modeling (from either ABS or investment casting wax materials). The surfaces of the patterns are then chemically treated to activate their surfaces to increase nucleation sites and to enhance the rate and quality of the mineral deposition. The scaffold pattern is then submerged in a simulated body fluid solution that deposits layers of biocompatible minerals on its surface.

The scaffold pattern is then submerged in a simulated body fluid solution that deposits layers of biocompatible minerals on its surface. Finally, the mineralized scaffold pattern is pyrolyzed to remove the pattern material and to sinter the deposited minerals, thus leaving behind a three-dimensional bioceramic scaffold.Critical to the development of this indirect scaffold fabrication process was the identification of mineralization processing parameters that would significantly increase mineral deposition. Using ANOVA, surface treatment and flow conditions were identified as significant mineralization factors compared to pattern material type. Plasma surface treatment and dynamic flow conditions were found to be ideal processing conditions and resulted in increased mineralization.The internal architecture of the fabricated scaffolds was characterized by the wall thickness (WT) and channel diameter (CD) and was benchmarked against existing indirect scaffold fabrication processes.

Although the smallest CD achieved was limited by the road width of the FDM process (~280 m), it is slightly better than the 300 m pores created by a direct inkjet printing process. The wall thicknesses achieved by the process were as small as 158 m – ~30% smaller than that achieved by existing indirect approaches. This fine feature size suggests that the biomimetic mineralization coating approach, which uses a low-viscosity aqueous SBF solution, alleviates the geometry and feature-size limitations found in existing indirect techniques that rely on viscous ceramic slurries.In future work, the authors look to perform cell viability and mechanical testing of the scaffolds. Qualitatively, the mineral structure seems too weak for applications other than maxillofacial implantation. In addition, the authors will investigate creating patterns with a biocompatible polymer, such as PCL, to mitigate concerns of ash residue from the polymers studied here.

In addition, the authors will investigate creating patterns with a biocompatible polymer, such as PCL, to mitigate concerns of ash residue from the polymers studied here. Furthermore, the authors will investigate the use of a direct inkjet printing AM process, such as the Objet PolyJet process, as a means for fabricating scaffolds with designed mesostructure. Such a process would alleviate the geometric limitations found in extrusion processes, which are limited to simple overlapping of parallel roads (aka “log piles”) [79] with a fixed height between layers [80].9Finally, the authors would like to further explore the crystalline surface structure that results from the mineralization approach. It is believed that there exist nano-sized topographical features that might influence cell adhesion and orientation. In addition, the resulting scaffold surface will likely provide micro-pores in addition to the designed macro-pores.

In addition, the resulting scaffold surface will likely provide micro-pores in addition to the designed macro-pores. This gradient porosity, which will be created without particle leaching, has been shown to be beneficial as larger pores increase nutrient diffusion while the smaller pores provide living spaces for cells [55].7. ACKNOWLEDGEMENTSThe authors acknowledge Dr. Aaron Goldstein, of the Virginia Tech Department of Chemical Engineering, for his suggestions and insight in the biomimetic mineralization process. We also acknowledge the Institute for Critical Technologies and Applied Science (ICTAS) Nanoscale Characterization and Fabrication Laboratory (NCFL) at Virginia Tech for support with FE-SEM imaging.8. REFERENCES1.2.3.4.5.6.7.8.9.10.11.12.13.14.15.16.17.Hutmacher, D.W., Scaffolds in tissue engineering bone and cartilage. Biomaterials, 2000. 21(24): p. 2529- 2543.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperIndividual and Coupled Contributions of Laser Power and Scanning Speed towards Process-Induced Porosity in Selective Laser MeltingSubin Shrestha, Thomas Starr, Kevin ChouAdditive Manufacturing Research Center University of Louisville Louisville, KY 40292AbstractPorosity is an undesirable characteristic of selective laser melting (SLM) process and keyhole pores are formed when the energy density is very high which leads to deep penetration melting. In this study, single-track SLM experiments using Ti-6Al-4V powder were designed and conducted with combination of varied levels of the laser power and the scanning speed, intended to obtain the same energy density. Three energy densities: 0.32 J/mm, 0.4 J/mm and 0.48 J/mm were selected to investigate the influence of laser power versus scanning speed on porosity.

Three energy densities: 0.32 J/mm, 0.4 J/mm and 0.48 J/mm were selected to investigate the influence of laser power versus scanning speed on porosity. Pore numbers and volumes was analyzed using micro-scale computed tomography. The results indicated that the pore formation is affected more by the change in the power than the scanning speed while keeping the energy density constant. As the power increased from around 20 W to 140 W, total pore volume increased, whereas pore volume decreased when power increased from 140 W to 195W.Keywords: Computed tomography; Porosity; Selective laser melting.1. IntroductionSelective laser melting (SLM) process has become a widely studied subject due to its freeform fabrication capability. However, inherent SLM defects such as porosity, part deformation are undesirable. Pores within the SLM fabricated samples have significant effect towards the part performance [1-3]. Leuders et al.

However, inherent SLM defects such as porosity, part deformation are undesirable. Pores within the SLM fabricated samples have significant effect towards the part performance [1-3]. Leuders et al. demonstrated that the fatigue strength can be optimized by reducing the pore size as the stress concentration is reduced [4]. Besides the size of pore, its location is also critical to the stress concentration: pores near to surface were found more critical even if the size were smaller [5].Pores, in general can be categorized as keyhole pores and metallurgical pores [6]. Keyhole pores are the result of keyhole mode melting which occurs when the energy density exceeds the keyhole threshold [7]. The energy density is the function of laser power, scanning speed, layer thickness and hatch spacing. Gong et al. [8] studied the effect of laser power and scan speed on the defect formation.

The energy density is the function of laser power, scanning speed, layer thickness and hatch spacing. Gong et al. [8] studied the effect of laser power and scan speed on the defect formation. Based on part quality, the power-speed curve can be divided into three zones: keyhole melting, conduction melting and incomplete melting. Pores can be formed when the energy density is very high or very low. Small amount of pores formed due to higher energy input are harmless when present up to 1 vol.% whereas defect due to insufficient energy input have effect even if the amount is as low as 1 vol.% [9]. Similarly, Aboulkhair et al. [6] investigated the effect of different process parameters to find the parameter window which1400would produce high density parts from AlSi10Mg. The metallographic study showed that different pores such as keyhole pores and metallurgical pores may form depending upon the process parameters used.

The metallographic study showed that different pores such as keyhole pores and metallurgical pores may form depending upon the process parameters used. Besides, with AlSi10Mg, hydrogen porosity may form due to the presence of moisture in powder particle surface as well as dissolved hydrogen in the powder material. Ponnusamy et al. [10] performed statistical analysis on porosity of 17-4PH stainless steel. Effect of various process parameters such as laser power, defocus distance, layer thickness and build orientation were investigated which also helped to find optimum process parameters. Although, metallographic studies clearly reveal the pores present inside the sample, it does not provide complete information of pore distribution inside the whole sample.As pore volume as well as pore location has significant effect on part performance, it is important to characterize pores formed due to different process conditions. Slotwinski et al.

Slotwinski et al. [11] performed porosity measurement on additively manufactured cobalt-chrome(Co-Cr) using X- Ray computed tomography (XRCT). XRCT technique help determine the morphology and distribution of pores within the sample. Siddique et al. [5] demonstrated that the percentage porosity calculated by two-dimensional metallography and three-dimensional tomography had no significant difference considering the critical pores. Kim et al. [12] utilized X-ray microtomography (µCT) to study the track changes in morphology of SLM Ti porous structures at different stage of post laser melting production. It has been demonstrated that µCT can be used not only for 3D quantification, but also as a feedback mechanism to make improvements on initial designs. Van Bael et al. [13] used XRCT as feedback mechanism to improve geometrical and mechanical controllability of selective laser melted Ti-6Al-4V porous structures. Ziolkowski et al.

Van Bael et al. [13] used XRCT as feedback mechanism to improve geometrical and mechanical controllability of selective laser melted Ti-6Al-4V porous structures. Ziolkowski et al. [14] explained the advantage of nondestructive XRCT over traditional metallograpic cross- sectional analysis. However, the major factor in XRCT is the resolution which would limit the pore size that can be studied.In this study, the individual and coupled effect of power and scanning speed on single track porosity has been investigated. Different power and scan speeds were used to obtain three energy densities: 0.32 J/mm, 0.4 J/mm and 0.48 J/mm. Single tracks samples were fabricated and the internal pore feature were studied with micro-CT. Pore count and pore volume resulting from different process parameter have been compared.2. ApproachSLM specimen fabricationsHollow cylindrical samples have been designed as shown in Figure 1(a). The single tracks are formed on top of the solid semi-cylinder base.

ApproachSLM specimen fabricationsHollow cylindrical samples have been designed as shown in Figure 1(a). The single tracks are formed on top of the solid semi-cylinder base. A constant spacing of 0.9 mm is maintained between single tracks and therefore each sample only contained four tracks. Two notches are used to identify the start and end of the single tracks. This design would also assist while performing CT scanning. Each sample has been marked as seen in Figure 1(a) to differentiate among the samples. In this study, three energy densities: 0.32 J/mm, 0.4 J/mm and 0.48 J/mm have been considered to study the effect of individual and coupled effect of power and scanning speed on porosity formation. The power and scanning speed used are presented in Table 1. The upper limit of power is maintained at maximum available laser power of 195 W. The design of experiment resulted in 12 tracks for 0.32 J/mm energy density while only 8 tracks are formed for14010.48 J/mm.

The design of experiment resulted in 12 tracks for 0.32 J/mm energy density while only 8 tracks are formed for14010.48 J/mm. Figure 1(b) shows the cylindrical samples with support on the build plate after fabrication. Three replicates have been produced for each sample.Figure 1: (a) CAD model of the specimen (unit in mm) and (b) Semi-hollow cylinders with single track built on the base plate.Table 1: Process parameters used to form single tracksED(J/mm) Speed(mm/s) 50 100 150 200 250 300 350 406.2 450 487.5 550 609.4 80 0.32 144 156 112 130 100 120 140 162.5 180 195 0.4 120 144 168 195 0.4816 32 20 40 24 4848 60 7264 80 9696176 195 - -Power (W)-CT scanningBruker SkyScan 1173 micro-CT scanner as shown in Figure 2 is used to perform CT scanning of the fabricated samples. To align the sample with the X-ray source, brass sample mount with appropriate size is used.

To align the sample with the X-ray source, brass sample mount with appropriate size is used. Clay, Styrofoam and parafilm tape were used to position sample on top of the brass sample mount as shown in Figure 2. Styrofoam and parafilm tape are preferred for the sample positioning as they are radiotransparent and do not affect the scanning of highly dense Ti-6Al-4V. Pixel resolution of 6 µm was used and samples were scanned at 2000 magnification. In order to transmit enough X-ray through highly dense titanium sample, high voltage of 130 kV x-ray source was generated which was filtered through a 0.25 mm brass filter so as to absorb x-ray with energy below 60 kV. This is necessary for highly dense materials to reduce beam hardening artifact otherwise the outer surface would appear more dense than interior due to higher absorbtion of low energy x-rays near outer surface.

After obtaining raw images, steps towards reconstruction such as beam hardening correction, ring artifact reduction, misalignment compensation and smoothing were performed using parameter fine-tuning to minimize the ring artifacts and blurring effects, etc [15].Figure 2: Specimen setup in the CT system (Left) and closed-up views (Right).14023. Results and DiscussionTypical scanning imagesFigure 3 shows example of CT images from a typical specimen with different sectional views as well as the isometric partial cutoff view. The color coding is based on the density of the material: light grey area is solid, medium grey area is powder and the dark spots in the YZ cross section, which represent density of air, are pores. Four single tracks formed atop of base pad are clearly identified. These four tracks, in the figure shown are formed with laser power and scanning speed of 120W and 250 mm/s, 144 W and 300 mm/s, 168 W and 350 mm/s and 195W and 406.2 mm/s from left to right.

These four tracks, in the figure shown are formed with laser power and scanning speed of 120W and 250 mm/s, 144 W and 300 mm/s, 168 W and 350 mm/s and 195W and 406.2 mm/s from left to right. The pores can be distinctly identified which are formed beneath the single track. These pores are mostly formed due to keyhole formation at high energy densities.Figure 3: Examples of CT images of a scanned SLM specimen.Single track morphologyCT scanning results were used to observe the track morphology as shown in Figure 4. It was noticed that at low power (24 W) and speed (50 mm/s), discontinuous track is observed even if high energy density is used. This may be due to lower power level: the input intensity is not enough to sufficiently melt the powder and substrate beneath to form a continuous melt pool. When both power and speed were doubled, a continuous track was observed. This track however had lower track width compared to other high power and speed cases.

When both power and speed were doubled, a continuous track was observed. This track however had lower track width compared to other high power and speed cases. Figure 5 shows the average single track widths obtained with different scanning parameters. The track width for 24 W and 50 mm/s has not been considered due to discontinuous track. 48 W laser power with 100 mm/s speed only resulted in average track width of 146 µm, whereas track width beyond this power and speed were over 190 µm. For same energy density, lower power and lower speed resulted in track morphology as that of low energy density cases. But the melt pool width did not1403increase with further increase in power and speed.

But the melt pool width did not1403increase with further increase in power and speed. This may be due to the keyhole mode melting: the flow is more dominant along the depth which resulted in almost similar melt pool width or track width.Figure 4: Single track morphology at different process parameters for ED=0.48 J/mm.Figure 5: Track width from different scanning parameters for ED=0.8 J/mm.Process effectThe transverse sections of single tracks have been presented in Figure 6. The transverse section of single track with 24 W, 50 mm/s shows wavy profile due to discontinuity in the track. As the power and speed were increased, the waviness was reduced and continuous, dense tracks were formed. Single tracks with 48 W and 72 W did not reveal pores, while as the power increased beyond 72 W, numerous pores were observed. This means that even for same energy density, keyhole mode melting highly depend on the level of power used.

This means that even for same energy density, keyhole mode melting highly depend on the level of power used. These pores are analyzed to compare the pore count and the pore volume among the fabricated single tracks.1404Figure 6: Transvere sections from single track with ED=0.48 J/mm.1405The pore analysis is performed over reconstructed image to obtain quantitative pore information. The raw images are transformed into binary images to calculate pore volume, surface area, sphericity etc. Then, 3D rendered pores are also generated. Figure 7 explains how the binary images of a pore at successive slices are used to obtain the rendered 3D representation of a pore.Figure 7: 3D rendering of a pore from 2D binary images [15].The effect of power and scan speed on pore count and pore volume have been summarized in Figure 8. As the power and speed increased for same energy density, the pore count and the total pore volume increased.

As the power and speed increased for same energy density, the pore count and the total pore volume increased. At low power and low speed, no or very few pores were observed. The pore count increased singnificantly around power 100 W and continued to increase until 140 W. In all three energy density cases, maximum volume of pores was observed when Power was around 140 W although scan speed were different: 144 W and 450 mm/s for 0.32 J/mm energy density, 140 W and 350 mm/s for 0.4 J/mm energy density and 144 W and 300 mm/s for 0.48 J/mm energy density. As the power and speed increased beyond this point, there is a significant drop in pore count and volume.Figure 8: Pore count and pore volume obtained at different process parameters from different energy densities (a) 0.32 J/mm, (b) 0.4 J/mm and (c) 0.48 J/mm.1406The penetration depth of keyhole is directly proportional to laser power and inversely proportional to the scan speed [16].

Also, power has more effect towards the maximum temperature in melt pool compared to scanning speed [7]. When we assume that melt pool depth normalized by the beam radius is only a function of ratio of deposited energy density to the enthalpy at melting:∆𝐻 ℎ𝑠=𝐴𝑃 𝜋ℎ𝑠√𝐷𝑢𝜎3Where, ∆𝐻 is specific enthalpy, ℎ𝑠 is enthalpy at melting, A is absorptivity, P is power, D isthermal diffusivity of the molten material, u is beam travel speed and σ is beam size.From equation 1, we may say that pore count and volume would increase with increasing power due to increase in keyhole effect. However this was not always true for the same energy density as increasing power after certain level decreased. It was interesting to see the similar trend in pore volume curve, with higher energy densities resulting in higher pore volume. We may also say that power is more significant towards the pore volume as the critical power seems to be similar for all three cases.

We may also say that power is more significant towards the pore volume as the critical power seems to be similar for all three cases. The relation 𝑃/√𝑢 is used to reduce the effect of speed on the graph which is shown in Figure 9. If magnitude of speed is further decreased, then the critical point would be aligned in a straight vertical line.From the track morphology and porosity results, we observe that from the experimental settings used, the critical power is around 140 W. When the power and speed is decreased from 140 W, the track width decreased along with the porosity. And when the power and speed increased, the track width remained almost constant while reduction in pore number and the pore volume. It would be interesting to see what would be the porosity level beyond power level of 195 W for same energy density.Figure 9: Coupled effect of power and speed on pore volumeThe pores are formed at different depth with different process parameters.

It is difficult to explain the penetration depth resulted from different power level with the study of porosity alone. However, we can make a statement that pore depth would provide initial idea on how deep the keyhole penetration is. The keyhole penetration seemed to increase with increasing power and speed. A metallographic study is required to have better understanding of this phenomenon.1407(1)4. ConclusionsPorosity analysis in SLM single-track specimens was performed using a SkyScan 1173 micro-CT scanner. The results clearly show pores resulted from the keyhole phenomenon and following conclusions can be made from this study: • The keyhole behavior is more affected by power than scan speed.

The results clearly show pores resulted from the keyhole phenomenon and following conclusions can be made from this study: • The keyhole behavior is more affected by power than scan speed. As the power increases forsame energy density, pore volume increased even if scanning speed increased.At same energy density, the lower power would result in balling behavior which is observed at low energy density melting: the input intensity may not be enough to properly melt the powder and substrate.Pore volume is maximum at around 140 W laser power. As the power further increased pore number and pore volume decreased for the interval of power investigated in this paper.The maximum power studied is limited to 195 W due to the limitation of EOS M270 power output. It would be interesting to see if the pore number and volume would further decrease with increasing power and scanning speed for studied energy densities.AcknowledgementsThis study is supported by U.S.

National Science Foundation under a grant (1662662).References[1] Cherry, J., Davies, H., Mehmood, S., Lavery, N., Brown, S., and Sienz, J., 2015, "Investigation into the effect of process parameters on microstructural and physical properties of 316L stainless steel parts by selective laser melting," The International Journal of Advanced Manufacturing Technology, 76(5-8), pp. 869-879. [2] Edwards, P., and Ramulu, M., 2014, "Fatigue performance evaluation of selective laser melted Ti–6Al–4V," Materials Science and Engineering: A, 598, pp. 327-337. [3] Siddique, S., Imran, M., Wycisk, E., Emmelmann, C., and Walther, F., 2015, "Influence of process-induced microstructure and imperfections on mechanical properties of AlSi12 processed by selective laser melting," Journal of Materials Processing Technology, 221, pp. 205-213.

205-213. [4] Leuders, S., Thöne, M., Riemer, A., Niendorf, T., Tröster, T., Richard, H., and Maier, H., 2013, "On the mechanical behaviour of titanium alloy TiAl6V4 manufactured by selective laser melting: Fatigue resistance and crack growth performance," International Journal of Fatigue, 48, pp. 300-307. [5] Siddique, S., Imran, M., Rauer, M., Kaloudis, M., Wycisk, E., Emmelmann, C., and Walther, F., 2015, "Computed tomography for characterization of fatigue performance of selective laser melted parts," Materials & Design, 83, pp. 661-669. [6] Aboulkhair, N. T., Everitt, N. M., Ashcroft, I., and Tuck, C., 2014, "Reducing porosity in AlSi10Mg parts processed by selective laser melting," Additive Manufacturing, 1, pp. 77-86. [7] King, W. E., Barth, H. D., Castillo, V. M., Gallegos, G. F., Gibbs, J. W., Hahn, D. E., Kamath, C., and Rubenchik, A.

77-86. [7] King, W. E., Barth, H. D., Castillo, V. M., Gallegos, G. F., Gibbs, J. W., Hahn, D. E., Kamath, C., and Rubenchik, A. M., 2014, "Observation of keyhole-mode laser melting in laser powder-bed fusion additive manufacturing," Journal of Materials Processing Technology, 214(12), pp. 2915-2925.1408[8] Gong, H., Rafi, K., Gu, H., Starr, T., and Stucker, B., 2014, "Analysis of defect generation in Ti–6Al–4V parts made using powder bed fusion additive manufacturing processes," Additive Manufacturing, 1, pp. 87-98. [9] Gong, H., Rafi, K., Gu, H., Ram, G. J., Starr, T., and Stucker, B., 2015, "Influence of defects on mechanical properties of Ti–6Al–4 V components produced by selective laser melting and electron beam melting," Materials & Design, 86, pp. 545-554. [10] Ponnusamy, P., Masood, S., Ruan, D., Palanisamy, S., and Mohamed, O., "Statistical analysis of porosity of 17-4PH alloy processed by selective laser melting," Proc.

545-554. [10] Ponnusamy, P., Masood, S., Ruan, D., Palanisamy, S., and Mohamed, O., "Statistical analysis of porosity of 17-4PH alloy processed by selective laser melting," Proc. IOP Conference Series: Materials Science and Engineering, IOP Publishing, p. 012001. [11] Slotwinski, J. A., Garboczi, E. J., and Hebenstreit, K. M., 2014, "Porosity measurements and analysis for metal additive manufacturing process control," Journal of research of the National Institute of Standards and Technology, 119, p. 494. [12] Kim, T. B., Yue, S., Zhang, Z., Jones, E., Jones, J. R., and Lee, P. D., 2014, "Additive manufactured porous titanium structures: Through-process quantification of pore and strut networks," Journal of Materials Processing Technology, 214(11), pp. 2706-2715.

D., 2014, "Additive manufactured porous titanium structures: Through-process quantification of pore and strut networks," Journal of Materials Processing Technology, 214(11), pp. 2706-2715. [13] Van Bael, S., Kerckhofs, G., Moesen, M., Pyka, G., Schrooten, J., and Kruth, J.-P., 2011, "Micro-CT-based improvement of geometrical and mechanical controllability of selective laser melted Ti6Al4V porous structures," Materials Science and Engineering: A, 528(24), pp. 7423- 7431. [14] Ziółkowski, G., Chlebus, E., Szymczyk, P., and Kurzac, J., 2014, "Application of X-ray CT method for discontinuity and porosity detection in 316L stainless steel parts produced with SLM technology," Archives of Civil and Mechanical Engineering, 14(4), pp. 608-614. [15] Shrestha, S., Shrestha, S., and Chou, K., "Porosity Analysis in Metal Additive Manufacturing by Micro-CT," Proc. ASME IMECE Conference(accepted).

608-614. [15] Shrestha, S., Shrestha, S., and Chou, K., "Porosity Analysis in Metal Additive Manufacturing by Micro-CT," Proc. ASME IMECE Conference(accepted). [16] Fabbro, R., 2010, "Melt pool and keyhole behaviour analysis for deep penetration laser welding," Journal of Physics D: Applied Physics, 43(44), p. 445501.1409.

viewed PaperINDUCTION INITIATED CURING OF ADDITIVELY MANUFACTURED THERMOSET COMPOSITESL. Omer*, M. J. Uddin*, H. Chowdhury*, J. Martinez*, I. Sporn*, B. Dudek* and J. Tate*†Ingram School of Engineering † Materials Science Engineering and Commercialization Program †Email: JT31@txstate.edu Phone: (512) 245-1826ABSTRACTAdditive Manufacturing (AM) has provided a new potential for material customization through the reduction in geometric constraints of products. The freeform geometry produced using AM allows for the creation of optimized structural parts that reduce weight and material waste. In addition, AM reduces production time by providing a model to product workflow with limited requirements for tooling and machine setup in advance. However, current work in printing thermoset composite parts is limited. One of the confounding factors in the AM of thermoset composites involves in situ curing of the composite resin system.

However, current work in printing thermoset composite parts is limited. One of the confounding factors in the AM of thermoset composites involves in situ curing of the composite resin system. This research proposes the creation of an additive manufacturing process for thermoset composites based upon induction initiated thermoset curing. The use of induction-based curing will allow the production of parts without the normal constraint of either oven or autoclave curing. A rapid curing resin system will be reinforced using milled fiber reinforcement and doped with a ferromagnetic susceptor to induce homogenous heating and curing of the resin in situ. The resulting composite will be evaluated for mechanical performance versus oven cure samples. Scanning electron microscopy in tandem with energy dispersive x-ray spectroscopy will be used to validate homogenous particle dispersion.

Scanning electron microscopy in tandem with energy dispersive x-ray spectroscopy will be used to validate homogenous particle dispersion. Lastly, differential scanning calorimetry will be used to compare the degree of polymer conversion between induction and conventional curing.Keywords: Fused Filament Fabrication, Thermoset Curing, Advanced Composites, Additive Manufacturing, Polymer Matrix Composites, Induction Heating1. INTRODUCTIONThe design and fabrication of lightweight composite materials are gaining popularity because of their enormous potential to replace traditional structural materials. The fact that weight reduction may be increased with topology optimization without sacrificing structural performance has sparked renewed interest in additive manufacturing (AM). Material production with complicated geometries is possible with this method.

Material production with complicated geometries is possible with this method. As a result of this versatile style of manufacturing, components can be altered, and their weights can be reduced without compromising the component's functionality. Continuous fiber-reinforced polymer composites headline the potential in this area since their mechanical properties are comparable to the metals, having significantly lower density [1].However, there are physical and processing limits in additive manufacturing of polymer composites reinforced with continuous fibers, including a minimum deposition length and minimal corner radius [2]. Manufacturing with short fibers provides much more flexibility in fiber2positioning and material deposition, resulting in easier material processing.

Manufacturing with short fibers provides much more flexibility in fiber2positioning and material deposition, resulting in easier material processing. The lower material costs and void content of short fiber composites, make them an appealing solution for many AM applications.The fused filament fabrication (FFF) additive manufacturing technology was utilized to develop short fiber-reinforced composites by melting and extruding a thermoplastic polymer matrix. According to Nawafleh’s Additive manufacturing of short fiber reinforced thermoset composites [1], the addition of short fibers to a plain, unreinforced thermoplastic matrix resulted in a considerable improvement in tensile strength and elastic modulus in several of these research studies. Although thermoplastic composites with significant quantities (40%) of short fibers could be produced and 3D-printed well enough, the maximum tensile strength of these composites was still limited (100 MPa).

This was attributed to poor interfacial adhesion between the fibers and the thermoplastic matrix, as well as porosity between the print lines (which was unavoidable in the FFF process). As demonstrated in Figure 1, these processing difficulties limited the strength and hence the applications of additively manufactured thermoplastic composites.Figure 1. Strength versus fiber volume fraction chart on the short fiber reinforced thermoplastic and thermoset composites. From Nawafleh’s AM study [1].The liquid resin in thermoset composites can moisten the fiber surface, facilitating the chemical bonding process. The direct write technique is used for liquid thermoset materials additive manufacturing of process instead of the FFF process, which is dependent on polymer melting and solidification. Viscous pastes are formed by combining liquid polymer resins with fiber reinforcements and rheology-modifying nanoparticles and then extruding them into the desired geometries.

Viscous pastes are formed by combining liquid polymer resins with fiber reinforcements and rheology-modifying nanoparticles and then extruding them into the desired geometries. Finally, the extruded material is cured into solid structures by heat or light. In 2014, Compton and Lewis presented direct write additive manufacturing to construct carbon-fiber reinforced epoxy composites, and it has since been used by other research groups to fabricate various thermoset reinforced composites with short carbon or Kevlar fibers [3]. When the same3amount of fiber loading is used, it is obvious from Fig. 1 that additively manufactured thermoset composites are stronger than thermoplastic composites.Additive manufactured short carbon-fiber reinforced composites have the potential to replace common structural metals like aluminum and low-strength steels, resulting in significant weight reduction.

The mechanical characteristics of these materials may be improved even more by fine-tuning the additive manufacturing settings and minimizing flaws caused by a large number of pores in the current specimens. Compression testing showed that removing these flaws will improve the mechanical performance of these composite structures substantially.In summary, due to the high amount of porosity and inadequate adhesion between the thermoplastic polymer and the fibers in these materials, short fiber reinforced thermoplastic composites are strength-limited, as illustrated in Figure 1. Because of the good chemical interaction between the fibers and the thermoset matrix, as well as their low porosity, thermoset composites have substantially greater strength.

Because of the good chemical interaction between the fibers and the thermoset matrix, as well as their low porosity, thermoset composites have substantially greater strength. However, additive fabrication of these systems is exceedingly challenging at 5% fiber loading, which is necessary to obtain significant levels of mechanical strength and stiffness.Typical processing of polymer matrix composite materials involves curing using heat transfer to get a complete cure. This is often done in convection ovens, by heating the surrounding air of a composite. It can also lead to agglomeration of nanoparticles that may be added for the purpose of increasing various properties of the composite, such as increased mechanical strength. The Army concluded that a process called induction heating may be a solution to this issue, as the manipulation of ferrimagnetic particles may be able to cure a composite in a more uniform manner, making the product more isotropic [4].

These magnetic materials, called susceptors, heat up when exposed to these electromagnetic fields. During this magnetization-demagnetization cyclic process, hysteresis losses occur if the material is magnetic, causing heating [4].Induction heating of ferrimagnetic nanoparticles in a self-healing thermoplastic adhesive is used to enable magnetic induction heating, while also enhancing the lap-shear strength of the adhesive. By utilizing the hysteresis loss mechanism, strong bonds between carbon fiber composites are achieved by repeatedly bonding and debonding the adhesive. One experiment achieved a bonding strength that was higher than any other thermoplastic based adhesives reported to date [5].The use of induction heating may be a solution to the curing issue, but due to the complex ratio of particles the mechanical properties of the samples were shown to have mixed changes in performance in Thomas Bayerl’s study on induction heating of thermoplastics [6].

When compared to neat samples, the tensile properties would decrease, while the Young’s modulus would increase across all filler ratios. The other property change was an increase in brittleness of the material leading to a 25% decrease in notched impact strength [6].Although electromagnetic induction offers high energy density and the ability to heat selectively and uniformly, this cannot be obtained without the use of susceptor fillers to convert electromagnetic energy into heat. Proper ratios of reinforcement filler to susceptor filler to matrix material with adequate working temperature must be further explored for optimum mechanical properties of the material [7].The motivation of our research is to determine a precise ratio of reinforcement fillers to magnetic susceptor to optimize curing time and maximize mechanical properties. We are primarily concerned with testing different loading levels of composites fillers using susceptor curing.

We are primarily concerned with testing different loading levels of composites fillers using susceptor curing. Utilizing additive manufacturing, hysteresis loss mechanisms for curing are required to reduce the necessity of expensive and large oven curing system.4When a high temperature material is required that incorporates additive manufacturing and induction heating, we need multiple reinforcement fillers to satisfy various needs. These reinforcement fillers will be evenly dispersed to facilitate released exothermic heat and help cure the surrounding thermoset material as it is being printed. In terms of the filler material, Fe3O4 is used as a susceptor to allow for induction curing, due to its ferrimagnetic properties. Milled carbon fiber material is added to the material to provide an overall increase in mechanical properties, and graphitized micro balloons are added to increase the strength-to-weight ratio.2.

Milled carbon fiber material is added to the material to provide an overall increase in mechanical properties, and graphitized micro balloons are added to increase the strength-to-weight ratio.2. MATERIAL SYSTEMSTo produce our thermoset composites, we dried reinforcement fillers at 70°C in oven for 2 hours to reduce residual moisture and maintain resin temperature during mixing. We UHTR-9863S resin as our base matrix which was placed in a 70°C water bath for 30 minutes to ensure a workable resin viscosity. Once heated, Titanium Butoxide catalyst was added at 5 wt.% and evenly dispersed using hand mixing for two minutes followed by two minutes of high-speed mechanical mixing. Following catalyst dispersion, the preheated fillers were steadily mixed into the catalyzed resin and then mechanically mixed at 650 rpm for two minutes. After mixing, the mixed composite pastes were printed using Lutum V4 to produce samples of each material composition with dimensions of 12.5mm x 130mm x 6mm.

After mixing, the mixed composite pastes were printed using Lutum V4 to produce samples of each material composition with dimensions of 12.5mm x 130mm x 6mm. The coupons were removed from the printer and set in the induction coil to cure the sample using induction curing. The coupons were also set in a convection oven at 250C for 2.5 hours to produce oven cured sample. The samples were then tested for porosity, tensile testing, particle aggregation.For this project eight treatment groups were initially present, these groups were compared by the ratios of their filler materials in high or low concentrations as seen in Table 1. The filler material consisted of Fe3O4, Milled Carbon Fiber (MCF), and Graphitized micro balloons (GMB). Due to poor heat performance the low concentrations of Fe3O4 (--+, -+-, -++, and ---), the samples were not tested for their mechanical properties or morphology.Figure 2.

Due to poor heat performance the low concentrations of Fe3O4 (--+, -+-, -++, and ---), the samples were not tested for their mechanical properties or morphology.Figure 2. Treatment group samples (+--, +-+, +++, and ++-)The four samples that were kept for experimentation were named in respect to the concentration value of the filler used. If a high concentration was used, it was denoted as + and if5low concentration was used it was denoted as -, and each followed the convention of Fe3O4 then MCF then GMB. The treatment groups were as follows: +--, ++-, +-+, and +++, and the concentration levels of each treatment group can be seen in detail in Table 1 below.Table 1.

The treatment groups were as follows: +--, ++-, +-+, and +++, and the concentration levels of each treatment group can be seen in detail in Table 1 below.Table 1. Concentration levels of Filler Material in parts per hundred of UHTM resinTreatment Group +-- ++- +-+ +++Fe3O4 Concentration MCF Concentration 30pph (High) 30pph (High) 30pph (High) 30pph (High)20pph (Low) 30pph (High) 20pph (Low) 30pph (High)GMB Concentration 15pph (Low) 15pph (Low) 25pph (High) 25pph (High)Following optimization of composite reinforcements, the effects of curing strategy on particle agglomeration was studied in the material composite yielding the highest performance. Identical composites of the (+++) composite were cured using either the previously mentioned oven curing or induction curing. Induction cure was performed using a 10 turn, 4-inch diameter solenoid coil mounted on a MTI corporation SP-25A heater for 9 minutes with a 50 KHz frequency and 0.165 Tesla magnetic field strength.

Induction cure was performed using a 10 turn, 4-inch diameter solenoid coil mounted on a MTI corporation SP-25A heater for 9 minutes with a 50 KHz frequency and 0.165 Tesla magnetic field strength. Following curing, the materials were examined for particle degradation and dispersion within the resulting composites.3. EXPERIMENTATION3.1 TGA and DTGAFor each treatment group the thermal analysis was performed on the SDT650, which would perform both the DTGA and TGA tests simultaneously, using the same parameters. The samples were each heated up to 1000°C at a rate of 20°C per min while using 50mL of Nitrogen gas per minute to isolate the environment.

The samples were each heated up to 1000°C at a rate of 20°C per min while using 50mL of Nitrogen gas per minute to isolate the environment. The results of which would then be plotted as a graph of both weight% vs temperature and the first derivative of the weight% rate of change over time.3.2 Mechanical TestingTo test the mechanical properties of the coupons we printed, we used the MTS Exceed electromechanical testing machine, and adhered to ASTM D638 to conduct our experiment properly. We took multiple measurements of the length and width of the cross-sectional area of each specimen, measured the gage length for each specimen, and recorded the crosshead displacement in millimeters and ramped loading in kilo Newtons with respect to time. We conducted 3 runs for 4 different specimens of each treatment group with various concentrations of fillers (+--, ++-, +-+, and +++).

We conducted 3 runs for 4 different specimens of each treatment group with various concentrations of fillers (+--, ++-, +-+, and +++). After completing these runs, we calculated the stress and strain of each run, and graphed them as shown in Figures 4-6. It is important to note that because the strain was estimated as a quotient of the crosshead displacement over the gage length. This is not representative of the actual strain experienced at the point of failure in the test coupon but is used as an accurate estimation of strain for analysis purposes.63.3 Hirox Digital Microscope ImagingThe treatment group samples (+--, ++-, +-+, and +++) were then checked for pore occurrence and pore size (μm2) by utilizing 3D Imaging techniques with low Range High Resolution Zoom Lens (MXB-2016Z) having the non-contact adapter (AD- 2016H) and Hirox RH-2000 software.

For measuring pore occurrence and pore size of the samples, 120X magnification of the Zoom Lens was used.3.4 Scanning Electron Microscope (SEM)A scanning electron microscope (SEM) is an instrument for imaging topography and for obtaining material information of conductive specimens using a focused beam of high-energy electrons. The samples were coated with gold (Au) by using an Imaging Sputter Coater and with colloidal graphite for earthing the samples before inserting the samples for morphological analysis. The samples were then analyzed with a magnification of 80x, 200x, 1000x, and 2,000x. The SEM provided a detailed orientation of the dispersion and sizes of MCF and GMB throughout the sample. Energy Dispersive X-ray Spectroscopy (EDAX) was utilized to determine the molecular makeup of the sample. The EDAX showed if the MCF and GMB are dispersed evenly throughout the sample or if it was clumped in singular locations. [8]4.

The EDAX showed if the MCF and GMB are dispersed evenly throughout the sample or if it was clumped in singular locations. [8]4. RESULTS AND DISCUSSION4.1 TGA and DTGAEach of the samples had similar shaped results from the TGA analysis and resulted in a remaining char yield ranging from ~81% to 86%. In Figure. 3. below, the different ranges in which weight is lost are outlined from the TGA by the blue lines. The first and second phases are where excess moisture, solvents and monomers are evaporated. The third phase is where the main portionFigure 3. DTGA of the four tested treatment groups. The weight loss graph shows the different ranges in which weight is lost and first derivative graph shows where the most weight loss occurs.7of decomposition occurs and where a majority weight percent is lost. The fourth and final phase is the burning step of carbon, i.e., the carbon MCF and GMB.

The fourth and final phase is the burning step of carbon, i.e., the carbon MCF and GMB. What is left afterwards is primarily ash residue and any remaining fillers that did not completely burn. In between phases four and five, the atmosphere within the TGA switches from pure nitrogen to regular air. This is to see what type of degradation would happen with oxidation added into the equation. In a pure Nitrogen environment, the TGA can monitor weight degradation entirely due to heat alone without the influence of oxidation.When comparing the four treatment groups char yields to the filler ratio combination, one can see that without a high concentration of GMB, the presence of MCF does not affect the overall char yield. However, if GMB is present in a high concentration with a low concentration of MCF, the char yield lowers. When all three fillers are used in high concentration, the resulting char yield is the highest of all samples tested.

When all three fillers are used in high concentration, the resulting char yield is the highest of all samples tested. This is due to the excellent thermal characteristics offered by the GMB and MCF. In Table 2, shown below, there is a detailed breakdown of each of the treatment groups. In the table, the maximum rate of decomposition for each sample was recorded to show how different compositions would affect the overall char yields and decomposition rates.Table 2.

In the table, the maximum rate of decomposition for each sample was recorded to show how different compositions would affect the overall char yields and decomposition rates.Table 2. Detailed analysis of DSC/TGA graphsTreatment Group1(+--) 2(++-) 3(+-+) 4(+++)Heat Flow Highest Peak (W/g) 5.272 7.982 7.156 6.567Highest Peak Temp (℃)607.1 684.6 633.7 597.1Heat Flow Lowest Peak (W/g) 0.925 1.873 1.966 1.066Lowest Peak Temp (℃)142.1 196.3 172.9 140.1Char Yield (%)83.92 83.93 82.93 86.37Heat Flow Reversal Point (℃) 262 337 309.6 247.64.2 Mechanical TestingAs shown from Figure 4, we determined that the Ultimate tensile strength of these samples was lowest using the ++- concentration of fillers, and highest using the +++ concentration of fillers. From this we can determine that the Milled Carbon Fibers (MCF) and Graphitized micro balloons (GMB) increased composite strength with increasing concentration.

From this we can determine that the Milled Carbon Fibers (MCF) and Graphitized micro balloons (GMB) increased composite strength with increasing concentration. These samples appear very linear, due to the large coefficients of determination (R2 values). These samples all remain in the brittle failure region, with small percent elongation. However, the two treatment groups that contained higher concentrations of GMB elongated more than the treatment groups with low GMB concentrations. Using statistical analysis, the concentration of MCF and GMB was found to significantly impact composite ultimate tensile stress and elongation.8) a P M(s s e r t S5 4.5 4 3.5 3 2.5 2 1.5 1 0.5 0Stress vs. Strain of Treatment Groupsy = 0.0102x R² = 0.972y = 0.0062x R² = 0.9959y = 0.0095x R² = 0.984y = 0.0068x R² = 0.99890.000.010.020.03Strain (mm/mm)0.04++- +-+ +++ +--0.05Figure 4. Stress vs.

Stress vs. Strain of treatment4.2.1 Statistical Analysis of Material ProprietiesThe graphs did not properly illustrate young’s modulus, Yield strength, or any necking. The main rise of the graph can be considered strain-hardening. Since these samples were able to be stretched to cause large strains, it resulted an undefined plastic range and did not follow Hooke’s stress vs. strain law in the elastic region. Ultimate Tensile Strength was determined to be the value right before fracture, and the data was tested with an ANOVA test. The difference in elongation, or measurement that captures the amount a material will plastically and elastically deform up to fracture, was also determine through an ANOVA test. ANOVA is a statistical approach that divides observed variance data into various components for further testing. For three or more groups of data, ANOVA can be used to learn about the association between the dependent and independent variables.

For three or more groups of data, ANOVA can be used to learn about the association between the dependent and independent variables. According to the test there is significant data to conclude that the Ultimate tensile strength and elongation changed with the addition of GMB and MCF. According to the data +++ resulted in the highest tensile strength and elongation followed by +-+, ++-, and finally +--. Table 3. below shows the mean, standard deviation, and std. Error of ultimate tensile strength for ++-, +-+, and +++. Table 4. Shows that the significant value in the difference of ultimate tensile strength is .002 which is less than .05, resulting in the conclusion that there is significant evidence to prove there is difference between the materials' tensile strength. Table 5. below shows the mean, standard deviation, and std. Error of elongation for ++-, +-+, and +++. Table 6.

Hirox Digital Microscope image of (+--) groupFigure 6. Hirox Digital Microscope image of (++-) groupFigure 7. Hirox Digital Microscope images of (+-+) groupFigure 8. Hirox Digital Microscope images of (+++) groupTable 7. Pore measurement of treatment group samples (+--, ++-, +-+, and +++)Parameter Average Pore Size (μm2) % Pore Coverage (μm2) Total Pore Occurrence Range of Pore Size (μm2)+--41137.5011.21%12126818.60++-33192.876.97%11149960.60+-+13941.998.29%2239000.79+++9694.485.50%2554130.10From Table 7, it is evident that the average pore size is higher in +-- group, which means the highest pore coverage of 11.21% in this group. However, the most pore occurrences are found in +++ group whereas the range of pore size is higher in ++- group.114.4 Scanning Electron Microscope ImagingThe morphological analysis of milled carbon fiber (MCF) was investigated with the help of FEI-SEM.

In Figure 9, the SEM micrographs show that there is minimal damage to the GMB following processing and curing. There does not appear to be major agglomerations occurring, supporting that the mechanical mixing was sufficient to ensure particle dispersion.abaFigure 9. SEM Micrographs a. Induction cured Milled Carbon Fiber (MCF) (80x) b. Damaged micro balloons (200x)bFigure. 10. SEM Micrograph of MCF Agglomeration of: a. Induction cured sample b. Oven cured sampleFigure 10a and 10b show MCF agglomeration in induction curing as compared to the MCF agglomeration in over curing. This indicates that due to the presence of Fe3O4 which enables the induction curing and ensures the high agglomeration and high dispersion of MCF, GMB. Hence,12increases the mechanical properties i.e. lowers the porosity and pore size compared to the oven cured samples. We investigated the aspect ratio of microfiber both in induction cured sample and oven cured sample.

lowers the porosity and pore size compared to the oven cured samples. We investigated the aspect ratio of microfiber both in induction cured sample and oven cured sample. Figure 11a and 11b show that there is no significant difference in terms of aspect ratio during induction curing and oven curing. This indicates that carbon fiber did not get damaged due to induction curing resulting in the same material properties as of oven cured sample.baFigure 11. SEM Micrograph showing Aspect Ratio of Milled Carbon Fiber of: a. Induction cured sample b. Oven cured sampleSimilarly, we analyzed both induction cured and oven cured samples in energy-dispersive X-ray spectroscopy (EDAX) to characterize the constituent's aggregation in the samples. Figures 12 and 13 show the live mapping of induction cured sample and oven cured sample. The mapping analysis shows that the MCF is evenly dispersed.abFigure 12. Live mapping using EDAX Map Analysis of: a. Induction cured sample b.

The mapping analysis shows that the MCF is evenly dispersed.abFigure 12. Live mapping using EDAX Map Analysis of: a. Induction cured sample b. Oven cured sample13The EDAX map analysis in Figure 13, shows that during induction curing the iron particles are evenly dispersed across and there is no significant agglomeration of iron. This indicates that mechanical mixing was sufficient for the dispersion of the iron particles across the resin matrix. Hence, these evenly dispersed iron particles helped to optimize the utilization of hysteresis heat during the induction curing process.abFigure 13. EDAX map imaging of Iron Particle dispersion of: a. Induction Cured sample b. Oven Cured sample5. CONCLUSIONThe growing use of lightweight composite materials in additive manufacturing has paved the way for our research to study the impact of induction curing on additively manufactured thermoset composites.

Considering the limitation of printing thermoset composite parts, this research has determined a precise ratio of reinforcement fillers to magnetic susceptor for optimizing curing time and maximizing the mechanical properties. The primary focus was associated with testing different loading levels of composites fillers using susceptor curing. For exploring such high temperature material that incorporates additive manufacturing and induction heating, multiple reinforcing fillers were required. These reinforcing fillers were evenly dispersed to promote the release of exothermic heat and aid in the curing of the surrounding thermoset material while printing. Due to its ferrimagnetic characteristics, Fe3O4 was utilized as a susceptor in the filler material to initiate induction curing. Milled carbon fiber material was added to the composite to improve overall mechanical characteristics and graphitized micro balloons were added to improve the strength-to- weight ratio.

Milled carbon fiber material was added to the composite to improve overall mechanical characteristics and graphitized micro balloons were added to improve the strength-to- weight ratio. Finally, the performance measurements of the resulting induction cured composite samples were evaluated and demonstrated in comparison with the oven-cured composite samples. Thermogravimetric analysis was performed and showed that the highest remaining char yield was allowed due to high concentrations of each filler material, and most of the weight percent loss occurred 450°C to 600°C. The SEM micrograph showed that the MCFs were dispersed evenly in the resin matrix without damaging the fibers. EDAX live mapping indicated that the iron particles were evenly dispersed resulting in better induction capacity of the samples for the induction curing.146.

EDAX live mapping indicated that the iron particles were evenly dispersed resulting in better induction capacity of the samples for the induction curing.146. ACKNOWLEDGEMENTSThe authors would like to thank every member of Analysis Research Service Center (ARSC) and Advanced Composites Laboratory of Texas State University for their support during the experiments of this research. We would also like to appreciate the help of Dr. Jitendra S. Tate for providing continuous guidelines towards our research objective.7. REFERENCES1. Nawafleh, N. and E. Celik, Additive manufacturing of short fiber reinforced thermoset composites with unprecedented mechanical performance. Additive Manufacturing, 2020. 33: p. 101109. 2. Blok, L.G., et al., An investigation into 3D printing of fibre reinforced thermoplastic composites. Additive Manufacturing, 2018. 22: p. 176-186. 3. Compton, B.G. and J.A. Lewis, 3D‐printing of lightweight cellular composites. Advanced materials, 2014. 26(34): p.

The components can be generated either firmly attached to a substrate or fixed in an easily separable mode. If supporting structures are employed, undercuts up to 90° are feasible, without, a process parameter dependent maximum angles of undercut below 90° are obtained.The process has been introduced into the market, labeled microSINTERING by3D-Micromac AG.1. Introduction:As mechanical engineering industry encounters a growing demand of µm-sized or µm- structured components for an increasing range of applications, miniaturization is presently rank- ing among the most important goals in product and tool development. Compared to still higher resolving techniques, selective laser sintering (SLS) bears the advantages of relatively low pro- duction cost and short processing times even for series productions of micro parts. Furthermore, next to prismatic or tapered microstructures, undercuts and hollows can be realized.

Furthermore, next to prismatic or tapered microstructures, undercuts and hollows can be realized. Manufactur- ers usually describe the performance of their devices in terms of accuracy or minimal layer thick- ness which yield but little information on the actual feasible minimal structural feature size. Until recently commercial metal SLS units were unable to generate micro parts smaller than 100 µm /1,2,3/.In the past five years research groups have worked on laser sintering of various metallic powders with pulsed or q-switched pulsed regimes /4,5,6/. Solid bodies were obtained. Wall thicknesses which were reported in two cases.

Solid bodies were obtained. Wall thicknesses which were reported in two cases. Considering these as a criteria for the resolution, it remained at values of 100µm and above.A novel modification of SLS commonly referred to as “laser micro sintering” together witha new setup, developed by Laser Institut Mittelsachsen e.V., extends the resolution to less than 30 µm for overall resolution, 20µm for ligaments and of 10µm for notches at aspect ratios of 12 and above, and presently a minimal roughness Ra of 1.5µm can be achieved with a minimum rough- ness of 1.5µm /7,8,9,10/.709Since its first public announcement in May 2003 laser micro sintering /7,8/ has gained ad- vertence and acknowledgement among toolmakers and users.

The once innovative set up - con- sisting of a hermetically closed sinter chamber and a special rake – which already fitted the needs for the handling and selective sintering of sub-µm grained metal powders - has been upgraded to higher efficiency and industrial applicability. The specific procedure including the laser sintering regime has become a sufficiently reproducible routine by which functional micro-freeforms are obtained from a number of powder materials. Ceramics laser micro sintering is still in develop- ment.2. Characteristics and Performance of the TechniqueA schematic of the process assembly is displayed in Fig.1.

Ceramics laser micro sintering is still in develop- ment.2. Characteristics and Performance of the TechniqueA schematic of the process assembly is displayed in Fig.1. The set-up consists of a sintering chamber, an attached turbo molecular vacuum pump, a ScanLab beam scanner with a scan field of 25x25mm2, a Q-switched Nd:YAG – laser (λ = 1064nm) with an output of 0.1-10W in TEM00 mode and 0.5-100kHz pulse frequencies, gate valves for various shielding and reaction gases as well as the power supply and the control unit for the coater and piston drives.Figure 1: Schematic set-up for laser microsintering (“chamber type 2”)The sinter chamber is a vacuum tight stainless steel casket which is bisected into an upper and a lower compartment. The bisection allows the drives and the controls in the lower compart- ment to be kept unaffected by the reaction atmosphere.

The bisection allows the drives and the controls in the lower compart- ment to be kept unaffected by the reaction atmosphere. The lid on top of the casket holds an inte- grated quartz glass window with transmission for the applied laser radiation.710The process platform has one or more cylindrical bores for the powder supply and one for thesample piston. Two or more rakes [Fig. 2] sweep the powder materials in a circular motion onto the sample piston (the sintering platform). This technique allows realizing vertical gradients of material or respectively grain sizes along the vertical axis of a 3D-micro body.The blades of the rakes are metal cylinders with a sharpened edge. Because of their geometry the rakes also serve as intermediate powder reservoirs [Fig. 3]. The pistons are tight for powders and liquids, which allows to process also emulsions and ceramic slurries.

Because of their geometry the rakes also serve as intermediate powder reservoirs [Fig. 3]. The pistons are tight for powders and liquids, which allows to process also emulsions and ceramic slurries. The chamber can be evacuated by the attached turbo molecular pump down to pressures of 10-3Pa and it can be charged with shielding gases or reaction gases at any pressure in the range between 10-3Pa up to 4x105Pa. Flushing with reaction gases, which is possible at pressures of ≥1Pa, makes the device applicable for laser chemical vapor deposition (Laser CVD).The 3D-micro bodies are generated by sintering the body cross section of the respective layerwith a special regime of q-switched laser pulses. The proprietary software IVS STL Converter (Version 1.0) that was developed especially for this purpose controls the sinter process. STL – data can be processed with a high resolution on a micrometer scale. Especially curves are exe- cuted at fast rates with high precision.

STL – data can be processed with a high resolution on a micrometer scale. Especially curves are exe- cuted at fast rates with high precision. Outline and filling parameters can be adjusted arbitrarily. This feature enables the operator to realize structural or density gradients in the micro body. Con- tinuously repeated calibration of the scanner is integrated into the software accounting for the fidelity and precision of the technique, even at high aspect ratios.Figure. 2: Multiple rakes and powder supplies allow varying compositionsFigure. 3: Ringblade serves as rake and pow- der storageWith single component powders the texture of the resulting solid area is not a closed coating of metal, but is more a network of craters or wedges that root about 5-10µm below the mean sur- face level with crests above between 1 and 3µm. Densities between 40% and 75% arise from those materials.

Densities between 40% and 75% arise from those materials. Blends, especially those consisting of a refractory and a lower melting metal, yield densities of 90% and above [1].711Fig. 4: Coiled tungsten ligamentFig. 5: Bihelical tungsten structureMetallic micro freeforms can be either fused to the metal substrate or attached to the substrate surface by narrow sinter necks, frail enough to be sheared off without destruction of the generated free form but stable enough to fix the part throughout the raking and sintering process.Fig. 6: Three nested hollow spheres.All the examples of functional tools presented further on in this report [Figs. 7-11] have been fixed to the platform in this separable mode during genera- tion. Description of the morphology of the interface between work piece and platform have been presented at earlier events /7,8,9,11/.The SEM views in Figs.

Description of the morphology of the interface between work piece and platform have been presented at earlier events /7,8,9,11/.The SEM views in Figs. 4-6 give an impression of the resolution of laser micro sintering and its efficiency in the realization of undercuts, which in the case of the presented samples were real- ized without the use of support structures. In Fig. 6, the base of the outer shell of the tree nested spheres starts with an angle of 30° to the platform surface which complies with an undercut angle of 70°.3.Functional Tools produced by Laser Micro sinteringThe functional parts displayed below are generated from a blend of metal powders. The toolsare employed as grip bits for micro manipulators or micro positioning devices. The drawings, which were released to our disposal by MiLaSys Technologies, show the functional environment of the components. Figs. 7 show a bit for the lifting and positioning of small lenses.

The drawings, which were released to our disposal by MiLaSys Technologies, show the functional environment of the components. Figs. 7 show a bit for the lifting and positioning of small lenses. The tool has an open and hollow cylindrical base. Via the three visible channels vacuum can be applied.712Fig 7a: Grip bit for the posi- tioning of an optical lensFig 7b: Blow-up of the func- tional environmentFig 7c: Lens positioning de- vice. Functional environmentThe wafer lifter bit in Figs. 8 has the shape of a square flat funnel and is employed to lift cir-cuit boards or wafers by aspiration.Fig 8a: Grip bit for a wafer lifter. The glare is from remain- ders of the cleaning fluidFig 8b: Blow-up of the Func- tional environmentFig 8c: Wafer lifter. Func- tional environmentFig 9a: Aspiration box of the fiber holderFig 9b: Grid plate for fiber aspirationFig 9c: Grip bits of tweezers713Figs.

Func- tional environmentFig 9a: Aspiration box of the fiber holderFig 9b: Grid plate for fiber aspirationFig 9c: Grip bits of tweezers713Figs. 9a-c show three parts of a fiber holder: A box (a), that holds a grid (detail view in b) onto which the fibers are attracted by aspiration through the channels in the box, and the two grips of tweezers by which the fibers are fixed mechani- cally (the tweezers in Fig. 9c are of a different type than the ones displayed in the drawing in Fig. 9d).A photograph view of a mechanical stop lock is presented in Fig. 10a together with a drawing presenting a cut open view of the construction (10c), which describes the openings and hollows of the part. The generation process had to be inter- rupted in the first assay because of a power breakdown in the beginning (Fig.

The generation process had to be inter- rupted in the first assay because of a power breakdown in the beginning (Fig. 10b), yielding the operator the chance of a photograph of the part at an unfinished stage.Fig 9d: Fiber holder deviceAnother component of the same machine with a similarpurpose is shown in Figs.11.Fig 10a:Stop lockFig 10a: Inside view of stop lock (unfinished part)Fig 10a: Stop lock; drawing of inside viewProcessing times for a product like the above pre- sented stop lock are listed below in the following chap- ter.Besides the above pre- sented applications, laser micro sintered tools have been tested amongst others as electrodes in EDM (electro- discharge machining) /11/. Although the generated elec- trodes principally functioned in the process, there is still need of development as far as material composition and process parameters are concerned.

Although the generated elec- trodes principally functioned in the process, there is still need of development as far as material composition and process parameters are concerned. With regards to material composition, multi- ple component selective laser sintering as described in chapter 2 could supply completely new possibilities.Fig 11a: Positioning ma- chine componentFig 11b: Inside view of machine component7144. Machining time effortThe stop lock presented in Fig. 10 was produced as a unicate in the chamber type 1 /7,8/ witha pulse frequency of 30 kHz. As it was the only part in the production cycle (batch size = 1) it required 3.6 hours to be built. The corresponding data set is outlined in the top section of Table 1. Depending on the type of laser, one can raise the pulse frequency considerably for the process.Table 1: Process times for the stop lock displayed in Figs.

10Chamber type 1 Batch sizeTotal coating time (in hours)Total sintering time (in hours)Total processing time (in hours)pulse frequency: 30kHz Process time per part (in hours)11.711.893.63.66 15 431.71 1.71 1.711.34 28.35 81.2713.05 30.06 82.982.18 2.00 1.93Chamber type 2 (upgraded coater) Total coating Batch time (in hours) size 0.61 1 0.61 6 0.61 15 0.61 43Total sintering time (in hours) 1.89 1.34 28.35 81.27Total processing time (in hours) 2.50 11.95 28.96 81.88pulse frequency: 30kHz Process time per part (in hours) 2.50 1.99 1.93 1.90Chamber type 2 (upgraded coater) Total coating Batch time (in hours) size 0.61 1 0.61 6 0.61 15 0.61 43Total sintering time (in hours) 0.95 5.70 14.25 40.85Total processing time (in hours) 1.56 6.31 14.86 41.6pulse frequency:75 kHz Process time per part (in hours) 1.56 1.05 0.99 0.96Chamber type 2 (upgraded coater) Total coating time (in hours) 0.61 0.61 0.61 0.61Batch size 1 6 15 43Total sintering time (in hours) 0.63 3.78 9.45 27.09Total processing time (in hours) 1.24 4.39 10.06 27.70pulse frequency: 100 kHz Process time per part (in hours) 1.24 0.73 0.67 0.64715We are presently working at a pulse frequency of 100 kHz and substrates (or platforms) which allow batch sizes of as many as 43 items equal or similar to the above mentioned part.

The difference between chamber type 1 and type 2 for the interpretation of Table 1 is merely the time required for the coating of the powder.Neglecting the loading and unloading of the chamber, the resulting machining time (“process time”) per part ranges between 40 and 45 minutes. Of course, as in all other versions of selective laser sintering, combination of different work pieces on a single platform is possible. Vertical stacking of several layers of work pieces, however, has not been performed yet.5.Suitability of the Sintered Material for Injection MoldsOne of the purposes of laser micro sintering is the generation of micro structured molds for pressure injection molding (PIM) or micro structured inserts into the molds.

The respective sin- tered material has to withstand the temperature changes and the occurring drag forces when the mold is opened and the product is ejected.Fig: 12a: Three protruding bars (2.5mm x 0.5mm) as a testing mold for the suitability of the material for PIM. The indentations are due to an inital false positioning of the countertool. Tool after casting experimentsTo prove the suitability of the sintered material a simple structure of three protruding horizontal bars was generated from an appropriate blend as an insert for a mold. The tool is displayed in Fig.12a (the in- dentations are due to false setting of the counter mold prior to the first cast). As demonstrated in Fig. 12b the tool allows the molding and de- forming of more than 500 casts with poly(oxymethylene) type Delrin 911P (“POM”).

As demonstrated in Fig. 12b the tool allows the molding and de- forming of more than 500 casts with poly(oxymethylene) type Delrin 911P (“POM”). As reported by our project partner, PORTEC GmbH, more subsequent casts from the same mold were performed with POM and other polymers under conventional and variotherm tem- perature regimes without any affect on the microstructure. The material proved dense enough to avoid infil- tration of the injected hot materialas well as sufficiently resistant to the drag forces during deforming.716Fig. 12b: The laser micro sintered mold remained unaffected after 500 casts with POMFig. 12c: POM-cast from laser micro sintered structure6. Conclusion and Perspectives:Since its introduction in May2003 laser micro sintering has gained advertence and acknowl- edgement among toolmakers and users.

Conclusion and Perspectives:Since its introduction in May2003 laser micro sintering has gained advertence and acknowl- edgement among toolmakers and users. The once innovative set up - consisting of a hermetically closed sinter chamber and a special rake – which already fitted the needs for the handling and selective sinter- ing of sub-µm grained metal pow- ders - has been upgraded to higher efficiency and industrial applica- bility. The specific procedure in- cluding the laser sintering regime has become a reproducible routine by which functional micro- freeforms are obtained from a number of powder materials.Figure 13: Prospective design of the commercial microSINTERING device which is built and will be mar- keted by 3D-Micromac AG, Chemnitz (Germany)The technique has proved suitable for the production of functional tools; material qualities can be achieved that meet the requirements for injection molds.

DUSTRIAL USE OF DIRECT METAL LASER SINTERINGOlli NyrhiHi, Juha Kotila, Jan-Erik Lind, Tatu Syvanen Electrolux Rapid Development Aholantie 17 FIN-21290 Rusko Finland Tel. +358-2-4365611 Fax. +358-2-4399620 Email: olli.nyrhila@notes.electrolux.fiAbstractThe Direct Metal Laser Sintering (DMLS) process was developed a few years ago by Electrolux Rapid Development (ERD) and EOS. DMLS has now been in commercial use for more than two years at Electrolux and also at other European prototype manufacturers, and the machine itself is also available for purchase.This paper describes the development of the process in particular the industrialization of DMLS and how it has changed the prototype mould manufacturing process at Electrolux. Issues concerning the materials used are also described, as well as the necessary post treatment methods in prototype mould manufacturing.

Issues concerning the materials used are also described, as well as the necessary post treatment methods in prototype mould manufacturing. Several case studies, from various injection mouldings to pressure die-casting, are presented.BackgroundDirect Metal Laser Sintering was commercially introduced in 1996. ERD received the first laboratory machine at the beginning of 1995. ERD had conducted the first laboratory tests even earlier than this, but a more professional development ofDMLS as an industrial manufacturing method started when the first laboratory machine was received.Some of the earlier research work on materials for DMLS has been reported previously in (1), however the history of the powder concept dates back to the end of the 1980's. This basic idea has been briefly described in (2), and more thoroughly in (3). Direct laser sintering of the proprietary bronze powder has been studied in one M. Sc. work (4).

This basic idea has been briefly described in (2), and more thoroughly in (3). Direct laser sintering of the proprietary bronze powder has been studied in one M. Sc. work (4). A more general survey of the laser sintering of steel powders is studied in one additional M. Sc. work (5).Apart from research and development work on materials used in direct laser sintering, the major task of the research team at ERD has been to develop new applications for laser sintered inserts. This work includes everything from surface finishing to mould concepts in the manufacture of injection moulds. The research group has been particularly successful in487being able to apply research and development results to in-house prototype production - one of the tasks set out for the group.DMLS Charasteristics2.1. Net-Shape OperationThe main goal of DMLS is net-shape operation.

Net-Shape OperationThe main goal of DMLS is net-shape operation. This aim was set, in fact, before rapid prototyping, because the sintering method which DMLS is based on is a net-shape furnace sintering process. There are several reasons why the net-shape process is beneficial and in some cases essential:- Accuracy -It is a simple operation: no shrink factors or compensations are neededThe potential to build large partsThe first argument is perhaps somewhat outdated, because of the ease in manipulating solids in computer models, but it is still an additional operation to take shrinkage into account in mould manufacturing. In layer manufacturing processes shrinkage is often different and z -directions and with free-formed surfaces this can create problems.Accuracy is probably the most important reason why net-shape operation is recommended.

Shrinkage almost never occurs uniformly and this is why all rapid prototyping methods that include shrinkage can meet the given tolerances only up to a certain length or thickness (6). Size limits in some rapid prototyping methods are due to this fact.The third argument can actually be derived from the second point. associated with the process, the only limitation for the part or mould size is the size of the manufacturing machine itself. In mould manufacturing with shrinking processes, apart from meeting the tolerances it is in many cases impossible to fit the mould halves together because distortions in different regions in the parting plane are not equal. If you cannot fit the mould halves together, the mould is useless.no shrinkage isAccuracy of the parting plane is naturally essential in mould making.

If you cannot fit the mould halves together, the mould is useless.no shrinkage isAccuracy of the parting plane is naturally essential in mould making. In a shrinking process it is often suggested that extra material be added to the parting planes, and then milled away so that the mould halves can be accurately fitted together. What is often forgotten, however, is that the same unpredictable amount of distortion is also taking place in the cavity of the mould, and if this is also not corrected the part is simply not usable. If both parting planes and cavities are machined, why use the free-forming method in the first place.2.2. Material PropertiesThe properties ofDMLS materials have been discussed in (2). One of the problems in dealing with porous materials is mechanical strength. Because DMLS is a net-shape sintering proce.ss, it inevitably produces porous material, which is naturally not as strong as fully dense matenal.

Because DMLS is a net-shape sintering proce.ss, it inevitably produces porous material, which is naturally not as strong as fully dense matenal. Development of materials has been done after reporting in (2). Normal tensile strength of the material near mould surfaces is about 200 MPa. The properties of the material depend488the x, ystrongly on the equipment used. The main factors are laser power and the quality of the laser beam.practice, weaker material has to be taken into account in the mould insert de~i~n. If a . designer thinks that a certain detail in a mould needs stronger material, the detmlln questIon simply has to be taken from a CAD file and made by other means. The detail is later added in the mould as a separate insert. This kind of hybrid construction is very common in prototype mould manufacturing, and normally does not cause serious problems.

The detail is later added in the mould as a separate insert. This kind of hybrid construction is very common in prototype mould manufacturing, and normally does not cause serious problems. It is also becoming more and more common in series mould manufacturing.Another possibility is to treat the critical regions as separate files. The detail files are then laser sintered together with the main part of the mould, but with extremely hard sintering parameters - essentially melting. As mechanical properties of the detail are then much greater than in greater and thus this approach can not be applied in cases of high volume. Combining normal direct laser sintering and melting in the detail dimension can also be preserved.mould insert generally, dimensional variations in the hard sintered detail areMechanical properties can also be enhanced by infiltration. Metal infiltration is normally increasing the density of sintered inserts.

Metal infiltration is normally increasing the density of sintered inserts. Unfortunately, this technique can only be applied used when dealing with materials that have a higher melting point, such as steel, because there are no suitable infiltrants available for bronze-based materials. There is also a risk of losing dimensional accuracy through the interactions between solid particles and the infiltrant. Mechanical properties gained after metal infiltration are comparable with the ones of the infiltrant metal itself. Normally in case ofDMLS, only polymer infiltration is used, more precisely, surface infiltration because the whole volume of the insert is not filled. The purpose of surface infiltration is to enhance surface properties, not to fill all pores of the insert.One problem associated with layer manufacturing is determining material properties. Before layer manufacturing the task was easier because everything was machined in standardized materials.

Before layer manufacturing the task was easier because everything was machined in standardized materials. In most layer technologies, the properties are different in the x, y and z -directions. DMLS the possibility of using different parameters in predetermined three-dimensionalsections of the body makes it even more difficult to characterize the properties in the conventional way.Surface PropertiesOne of the biggest problems in direct laser sintering methods, and also in DMLS, is the insert surface quality of the after sintering. Direct laser sintering is quite an aggressive method, a combination of sintering and melting. The sintered material, which is partly agglomerated particles, partly solidified melt pools, cannot form a very good surface directly from laser sintering. Therefore all critical surfaces have to be finished.A special polymer infiltration can be used to enhance surface finish.

Therefore all critical surfaces have to be finished.A special polymer infiltration can be used to enhance surface finish. This infiltration differs from conventional methods, because only pores on and just beneath the surface are filled. The method is called surface infiltration, and it uses a mixture of suitable metal particles and liquid polymer. Because the insert is porous, capillary forces drag the liquid inside the body. Particles move with the liquid deeper inside the pores and gradually close most of the pores on the surface.489After surface infiltration there are several possibilities for finishing the surface. Some prefer hand finishing, some try to apply automated NC-based methods. The quality of finish depends on the method chosen. Ifthe quality is expressed in R scale, values of about 0.5 Ra and about 2 Rmax can be reached by applying present methods. These values are subject to change because surface finishing is always undergoing intensive development.3.

These values are subject to change because surface finishing is always undergoing intensive development.3. Industrial UseThe use of DMLS inserts, as well all other mould inserts, differs considerably from the use of the other 'conventional' rapid prototyping parts. Service bureaus dealing with rapid prototyping machines normally just deliver rp parts. As input data they get STL files. In this kind of business the value added remains low. This can also be seen in the performance of many service bureaus - they are quite vulnerable to fluctuations in demand.3.1. Traditional Mould IndustryThe plastic component industry has been strictly divided into separate fields. Some companies manufacture tools, and some mould the plastic parts.

Traditional Mould IndustryThe plastic component industry has been strictly divided into separate fields. Some companies manufacture tools, and some mould the plastic parts. This means that the most of the industry is quite specialized, and knowledge of the entire development process is very rare.The development of the injection mould chain can be troublesome, because the process has to be done simultaneously in several different places. If a mould maker develops a new design concept, good results cannot be achieved unless the injection moulder is deeply involved in the development. All the problems are emphasized in the production of prototype molds, because of all additional trouble associated with product development. This difficulty has been seen many times in the history of laser sintered mould inserts.3.2. Industrial Use at ERDAt ERD the traditional way of making injection mold business was almost totally rejected years ago.

Industrial Use at ERDAt ERD the traditional way of making injection mold business was almost totally rejected years ago. This was done even before the era of rapid prototyping. ERD realized that to be effective in product development, one must be able to control the whole development chain, from component design to the manufacturing process, including all prototype-manufacturing processes.In prototype services, all rapid prototyping and rapid tooling methods are but a small part of the prototype production process. Services provided by an rp machine should fulfill their tasks in the development process. The tasks are many and they depend naturally on the project. This inevitably leads to the fact that there is not just one right method, not even for rapid tooling.Because ERD specializes in product development, DMLS is only one of many tools used handle a large variety of product development projects.

The reason why ERD developed the DMLS was to satisfY the internal need for such a machine. The same motivation led to the development of the furnace sintering process, which DMLS is based on.Normal DMLS project starts when a plastic component CAD-data is received. If the part is suitable for DMLS, a CAD-designer specialized in DMLS starts to design the mould. The490mould design covers both the insert and the mold itself. ERD's knowledge of the strenghts and weaknesses of the DMLS process is especially important when it comes to the design of the insert.The project manager makes the choice between different manufacturing methods. DMLS is suitable for moulds with extremely complex surfaces. It measures up very well with milling if large amounts of material have to be removed by this process. In cases where milling is impossible, DMLS is the best choice.

It measures up very well with milling if large amounts of material have to be removed by this process. In cases where milling is impossible, DMLS is the best choice. If spark erosion is going to be the manufacturing method for the production mould, DMLS is an extremely competitive alternative.The CAD designer provides the DMLS operator with a net-shape STL model. At this stage, the DMLS operator has numerous possibilities to alter the material properties at different regions of the insert. The parameters are adjusted so that strong material is produced at regions where maximum stress is expected. It is advantageous if the DMLS operator has knowledge of laser sintering materials. At ERD this is the case, because the laser sintering machines are operated by the DMLS materials development team.The goal in the production of DMLS inserts is to be able to do it then net-shape. At present, it can be done with an overall accuracy of ±O.05 mm.

At present, it can be done with an overall accuracy of ±O.05 mm. This accuracy is good enough in most of the injection moulding cases. The accuracy depends both on the material and also on the machine condition - particularly on the laser power. The interaction between the laser and the material varies with the power of the laser and this variation cannot be entirely compensated for by adjusting the scan speed.After laser sintering the insert undergoes surface infiltration, as described in chapter 2.3. The aim of surface infiltration is to make polishing easier by filling surface pores. Polishing is at the moment the most human labour-intensive process in the chain, and a lot of effort is being put into improving post-treatment methods. It seems, though, that it is extremely difficult to combine good surface quality and direct laser sintering. Some surface treatment is always needed.

It seems, though, that it is extremely difficult to combine good surface quality and direct laser sintering. Some surface treatment is always needed. A completely different question is whether this task can be automated or not.The typical turnaround time of a DMLS project from receiving the component's CAD-data to the delivery of the plastic parts is two weeks. If everything goes well, smaller components can be delivered in one week. Parts that are complicated and/or large can sometimes require three or four weeks. Figure 1 shows a typical injection mould and plastic parts, suitable for DMLS. The parting planes are complicated, and an alternative manufacturing method would probably be spark erosion, which is very time consuming. This mould was used as a small series manufacturing tool, and the number of plastic parts was in the thousands.491Figure 1 Typical DMLS Mould and Plastic PartsThe maximum mould size a DMLS machine can manufacture at present is 250 x 250 x 150 mm.

Larger moulds can be manufactured if the insert is divided into smaller sections the machine can handle. The injection mould is then assembled from these sections. The assembly is not very easy and it requires accurate manufacturing of the inserts. shows a large double injection mould. The maximum size is about one meter.2Figure 2 Large Double Injection Mould Assembled from SeveralofDMLSThe suitability of DMLS in various plastic and rubber moulding processes has been proven in many projects. The next major application field will be die-casting. done and the results hold promise.testing has been492materials development, the next step is stronger powder material. The goal, however, is toconsistently make net-shape inserts. Good results have been achieved with steel powders. Figure 3 shows a die-casting insert made from proprietary steel powder.Figure 3 Die-casting Insert Made In Steel1. M. Sindel,Pintat, M. Greul, O. NyrhiHi, C.

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperINFERENCE OF METAL ADDITIVE MANUFACTURING PROCESS STATES VIA DEEP LEARNING TECHNIQUESRichard Anarﬁα*, Benjamin Kwapongα, Kenneth Fletcherα, Todd Sparksβ, Aaron Floodβ, and Mugdha JoshiγαUniversity of Massachusetts Boston, Boston, MA 02125 βProduct Innovation and Engineering LLC, St. James, MO 65559 γMissouri University of Science and Technology, Rolla, MO 65409 *Corresponding authorAugust 20, 2021AbstractNumerical simulation of metal additive processes are computationally intensive tasks. Iterative solution techniques for physics-based methods can lead to lengthy solution times and convergence problems, particularly if ﬂuid dynamics of the melt pool are considered. Deep learning (DL) tech- niques offer an opportunity to infer solution results quickly.

Deep learning (DL) tech- niques offer an opportunity to infer solution results quickly. In this paper we propose a DL method based on long short term memory (LSTM), network trained on rendered images from a metal AM process simulation and CAM data. We obtained vector representations of the images by training on an autoencoder. LSTM is a memory based recurrent neural networks (RNN) that is capable of processing long sequences of data while combating temporal stability problems encountered with conventional recurrent neural networks (RNN)s. This LSTM network is used to predict images of the process given scan path and process information. This could later be used to compare with process monitoring systems as part of a quality assurance or process control schema. Key words: Additive manufacturing, autoencoder, deep learning1Introduction and MotivationAdditive manufacturing (AM) processes rely on a sets of technology parameters (e.g.

Key words: Additive manufacturing, autoencoder, deep learning1Introduction and MotivationAdditive manufacturing (AM) processes rely on a sets of technology parameters (e.g. slicing settings, inﬁll settings, speeds, feeds, laser settings) to generate functional machine level code to produce printed parts. Sets of the technology parameters are commonly referred to as a ‘proﬁle’. The contents of a particular proﬁle can be dependent upon process, AM system architecture, ma- terial, and desired part geometry. Complicating matters further, performance characteristics of the printed parts are a consideration in the proﬁle selection for metal parts produced via Directed Energy Deposition (DED) and Laser Powder Bed Fusion (LPBF) processes.Developing a new 3d print proﬁle can be a costly endeavor. Users typically rely on some combi- nation of experimentation and simulation to justify decisions and set values for critical parameters. This process can be both time consuming and costly.

Users typically rely on some combi- nation of experimentation and simulation to justify decisions and set values for critical parameters. This process can be both time consuming and costly. This paper presents an attempt to shortcut1295the costly computational methods used to simulate metal 3d printing processes by inference using a LSTM [1, 2]. When successful, this inference schema has applications in process development, machine controls, and quality assurance.2 Literature reviewSeveral researches have proposed different neural network architectures based on build classiﬁ- cation depending on the input laser process parameters. long short term memory (LSTM)-recurrent neural networks (RNN) are widely used in applications such as text completion [3, 4], image cap- tioning [5], stock prices forecasting [6]. LSTM networks can handle long term dependencies for long data sequences therefore are advantageous over traditional recurrent neural networks (RNN) architectures [7].

LSTM networks can handle long term dependencies for long data sequences therefore are advantageous over traditional recurrent neural networks (RNN) architectures [7]. Other than the mentioned applications, its applications in image series prediction is analogous to the current requirement. These networks can efﬁciently predict the trajectory of a pedestrian [8, 9, 10] as well as the trafﬁc trajectory [11, 12, 13].Different Neural network architectures have been used in the applications of in-situ monitor- ing and optimizing process parameters for DED processes[14, 15, 16]. However, recent studies show that this state of art technology is primarily used for prediction of temperature history in the DED process. According to the present literature, recurrent neural networks consisting of LSTM cells provide an accurate mapping of temporal information about thermal distribution in the DED process.A study by Ren et al.

[17] shows that in comparison of individual architectures, a combination of recurrent neural networks (RNN) and DNN shows consistency in contour maps and temperature distribution produced by ﬁnite element simulation. This type of architecture has an added advan- tage over individual networks that LSTM cell in recurrent neural networks (RNN) learns sequential data of laser scan path which is mapped to temperature ﬁeld information via fully connected layer in DNN. However this model is conﬁned to single layer process prediction with limited input information about laser process parameters and material properties.In a similar study Mozaffar et al. [18] proposed that a stacked recurrent neural networks (RNN) with Gated Recurrent Unit (GRU) formulation is efﬁcient in predicting thermal history at a single point for different geometries.

[18] proposed that a stacked recurrent neural networks (RNN) with Gated Recurrent Unit (GRU) formulation is efﬁcient in predicting thermal history at a single point for different geometries. GRU structure in the network can accurately ﬁnd out the hidden correlations in the high dimensional data and predicts critical features in the thermal history such as sharp gradients, and melting and re-melting of the material.Zhang, Liu, and Wu [19] used two Machine Learning (ML) algorithms XGBoost and LSTM- RNN to predict the melt pool temperatures of an eight layer thin-walled CarTech® 718 alloy sample made by DED process. Prediction accuracy of both algorithms was measured for different input process parameters. It is found out that LSTM outperforms XGBoost algorithm in all four special cases with highest and lowest values of Laser power and scan speed.

It is found out that LSTM outperforms XGBoost algorithm in all four special cases with highest and lowest values of Laser power and scan speed. Though LSTM is computationally ineffective than XGBoost, ﬂuctuation in the melt pool temperature does not affect the accuracy of LSTM model.Thus, LSTM model provides promising results in predicting thermal history.However, as far as we know, no previous research has been investigated in prediction of process images based on the laser scan path and process parameters. Although a lot of work has already been done in modeling and simulation of DED process, this technique is time consuming and cannot always guaranteed to be convergent. On the there hand, deep learning requires huge training data which is not easy1296to generate in AM processes.

On the there hand, deep learning requires huge training data which is not easy1296to generate in AM processes. Thus, this research proposes an innovative approach of integrating physics based iterative methods with deep learning techniques.3 Research methodologyIn this section, we describe the structure and various sections of our approach, starting with thegeneral overview, followed by detailed description of the sections/parts.Our model trains an autoencoder (Section 3.1) to learn a compressed (encoded) version of the image we feed into it. This is achieved using several convolutional layers (Section 3.3). The compressed image (from a vector of length >49,000 to a vector of length 16) can then be de- compressed (decoded) back into its original shape/dimension. Our model feeds the time-series embedded vectors of very small dimension (output of the autoencoder’s encoder) into an LSTM (Section 3.2) which we train to predict vectors that are some time steps ahead.

We then use our decoder from the autoencoder to decode these vectors into actual image vectors. The architecture of our model can be seen in Figure 1.3.1 AutoencoderThe concept of autoencoders as seen in Figure 2 is that we want to force a network to try to reconstruct our data and hopefully it will learn a useful representation of the data. For traditional autoencoders this is used for feature learning. Normally the encoder will be a 4-layer downsampledFigure 1: The architecture of our proposed model. There are 3 main parts: The encoder, which consists of convolutional networks, the LSTM, and the decoder which also has convolutional net- works.1297Figure 2: Autoencoderconvolution, and the decoder a 4-layer upsampled convolution. Sometimes they share weights: dim(x) = D, dim(z) = H, we : H × D, wd : D × H = wT e .The idea is to pass the input data through an encoder network that will produce features z. This step can be thought as PCA.

This step can be thought as PCA. We are transforming the input data and transforming it into another feature representation. The encoder network is usually a ReLU CNN. Normally z is smaller than x (dimensionality reduction). The problem is that we do not have any explicit labels to use to know which information is relevant and which one its not to decide how to make the dimensionality reduction. To decide which is the most relevant data we reconstruct the input data only with the information in the z features and compare the reconstruction with the original image. Specially if the z features are small, hopefully, it will force the network to summarize the useful statistics and to discover useful features of the input image.3.2 Long-short term memory unitsIn this section we describe the deep LSTM. We let subscripts denote timesteps and superscripts t ∈ Rn be a hidden state in layer l in denote layers. All our states are n-dimensional. Let hl timestep t.

The LSTM can decide to overwrite the memory cell, retrieve it, or keep it for the next timestep. The LSTM architecture used in our experiments is given by the following equations [20]:, hl i sigm f sigm o sigm g tanh cl t = f (cid:12) cl t−1 + i (cid:12) g t = o (cid:12) tanh(cl hl t)LSTM : hl−1 t   t−1, cl t−1 → hlt, cl t(cid:18)hl−1 t hl(cid:19)        =T2n,4nt−1In these equations, sigm and tanh are applied element-wise. Figure 3 illustrates the LSTM equa- tions.3.3 Convolutional Neural NetworksCNNs (Figure 4) generally consist of three elements: convolutional layers (Figure 5), pooling layers (Figure 6), and fully connected layers. In the convolutional layer, ﬁlters are convolved with the receptive ﬁeld of the input image in a sliding window style to learn data-speciﬁc features.

In the convolutional layer, ﬁlters are convolved with the receptive ﬁeld of the input image in a sliding window style to learn data-speciﬁc features. Basic features, such as lines, edges, and corners, are learned in the initial layers, while more abstract features are learned as layers go deeper [22].Generally, a pooling layer follows each convolutional layer. Max-pooling is basically a nonlin- ear down-sampling procedure, which takes the maximum of 2x2 neighborhoods of the image, and helps to reduce the computational complexity for the forward layers, as well as adding translation invariancy to the network. Fully connected layers are used to learn the nonlinear combinations of extracted features from previous layers. Dropout is recommended as a way of preventing overﬁt- ting by disabling randomly chosen neurons and their connections [23].

Dropout is recommended as a way of preventing overﬁt- ting by disabling randomly chosen neurons and their connections [23]. The dropped neurons stay inactive during the feedforward and backpropagation phases, thus forcing the network to learn different nonlinear combinations of features on each epoch.1299convolutional layer with non-linearities layer l = 1convolutional layer with non-linearities layer l = 4fully connected layer layer l = 7input image layer l = 0subsampling layer layer l = 3subsampling layer layer l = 6fully connected layer output layer l = 8Figure 4: The architecture of the original convolutional neural network, as introduced by LeCun et al. [24], alternates between convolutional layers including hyperbolic tangent non-linearities and subsampling layers. In this illustration, the convolutional layers already include non-linearities and, thus, a convolutional layer actually represents two layers.

If layer l is a pooling and subsampling layer and given m(l−1) = 4 feature maps of the previous layer, all feature maps are pooled and subsampled individually. Each unit in one of the m(l) 1 = 4 output feature maps represents the average or the maximum within a ﬁxed window of the corresponding feature map in layer (l − 1).14 Simulation DescriptionWith their work in AM [26, 27], Product Innovation and Engineering LLC (PINE) has devel- oped several physics-based simulations in order to better understand the AM process. The sim- ulation used in this study was developed to speciﬁcally target the thermal history of the material during a build. This simulation has the aims of being computationally efﬁcient while still being based on physics. It has the expressed goals of being fast and giving general guidance to path planning by giving a thermal history for a given build.

It has the expressed goals of being fast and giving general guidance to path planning by giving a thermal history for a given build. In order to accomplish this goal, a voxel based simulation was developed which heavily leverages a graphics processing unit (GPU) and is based on many image processing techniques. From past simulation development experience, it was understood that the calculation of the ﬂuid ﬂow is the most computationally expensive part of the simulation, therefore it was omitted in this simulation for the sake of efﬁciency.The simulation is able to predict the thermal history of a part, Figure 7a, the phase map of the part at any given time, Figure 7b, and a cooling rate for any section of the part, Figure 7c. This helps to give a predictor of the micro-structures within the deposition which are the driving force behind the ﬁnal mechanical properties.The laser in the simulation is modeled in 3-D to be able to take into account the beam quality.

In many lasers the beam quality is deﬁned using the beam parameter product (BPP) this is deﬁned as 0.5θw0 where θ and w0 are the divergence angle and the beam waist respectively. These are shown graphically in Figure 8a. When this is done in 3 dimensions the results can be seen in Figure 8b. In this ﬁgure, the laser proﬁle for each slice is a Gaussian proﬁle. In addition to taking into account the quality of the laser, it is ray traced onto the object in order to account for shadowing of the build by the incoming material and non-planar effects where the center axis of the laser is not perpendicular to the work surface.Other key features which have been developed in the system are the inclusion of temperature dependent material properties and the inclusion of work holding, shown in Figure 9. The larger section of material on the bottom of the substrate is the work holding.

The larger section of material on the bottom of the substrate is the work holding. This work holding has modiﬁed material properties so that it can emulate the mass and material properties of whatever1301(a) Temperature proﬁle(b) Phase map(c) Cooling rate mapFigure 7: Examples of data maps which can be expected from the simulation.(a) 2-D laser representation, where the black lines are the the laser, θ is the edges of divergence angle and w0 is the beam waist.(b) 3-D representation of the laser proﬁle.Figure 8: 2-D and 3-D representation of the laser proﬁle.1302work holding is used in the process.Figure 9: Example of the the inclusion of work holding in the simulation.5 Case StudyThe simulation that was described in Section 4 was used to generate the training data for theML model.

The parameters used in the simulation can be seen in Table 1.Table 1: Parameters used in simulationParameter Material Resolution Laser diameter Laser power Laser proﬁle Laser scan speed Powder stream diameter Powder stream rateValue Ti-64 250 µm 4.0 mm 1000 W TEM00 1273 mm/min 15 mm 12 g/minThe training images were created in an attempt to emulate images which could be collected from a physical camera on deposition systems. These images were generated using the colors of steel which can be seen in Figure 10. These colors were used because of the extensive character- ization which has been performed on steels and not been performed on Ti-64. These colors will serve as a starting point and proof of concept. In order to generate the images, the virtual camera was placed 50 mm above the melt pool and an image was created which was 1080 pixels by 1080 pixels based on what a physical camera would image.

In order to generate the images, the virtual camera was placed 50 mm above the melt pool and an image was created which was 1080 pixels by 1080 pixels based on what a physical camera would image. During the image generation, cold voxels which were recently inserted because of powder insertion were eliminated to clean up the image. Figure 11 shows a selection of the original and reconstructed (predicted) images from the test set after training of the autoencoder model.Figure 12 shows the results of the next image prediction task for a test set. Its corresponding graph of mean squared error and structural similarity score values are shown in Figure 13. This is the combined output of both the autoencoder and LSTM models.

Its corresponding graph of mean squared error and structural similarity score values are shown in Figure 13. This is the combined output of both the autoencoder and LSTM models. As can be seen from the graph, all1303Figure 10: Temperature to color map used in image[28]Figure 11: Image Reconstruction results from the trained autoencoder modelmean squared error values are below 0.001 across the test distribution and all structural similarity score values are above 0.985.It was interesting to observe how our model will behave in the event of a change in direction. Figure 14 shows the results of the next image prediction task for a test set which covers a change in direction range. Its corresponding graph of mean squared error and structural similarity score values are shown in Figure 15. This is the combined output of both the autoencoder and LSTM models.

Its corresponding graph of mean squared error and structural similarity score values are shown in Figure 15. This is the combined output of both the autoencoder and LSTM models. As can be seen from the graph, all mean squared error values are below 0.001 across the test distribution except for a tiny spike to 0.004 around timestep 5 before it stabilizes and all structural similarity score values are above 0.97 except for a tiny spike to 0.94 around timestep 5. This is consistent with what one can observe from the images in Figure 14 as a start to a change in direction at timestep 5.For an anomaly detection case study, we obtained simulated images with holes in the metal for the AM process at random timesteps. Figure 16 shows the results of the next image prediction task for a test set which covers examples of images with holes. Its corresponding graph of mean squared error and structural similarity score values are shown in Figure 17.

Its corresponding graph of mean squared error and structural similarity score values are shown in Figure 17. As can be seen from the graph, all mean squared error values are below 0.001 across the test distribution except for huge1304Figure 12: Autoencoder normal and predictedFigure 13: Autoencoder normal and predicted graphFigure 14: Autoencoder normal with change in directionFigure 15: Autoencoder graph of normal with change in directionspikes at timesteps 4 and 8 and all structural similarity score values are above 0.97 except for huge spikes at timesteps 4 and 8. This is also consistent with what one can clearly observe as holes in the images in Figure 16 at timesteps 4 and 8.1305Figure 16: Autoencoder normal vs anomalyFigure 17: Autoencoder normal vs anomaly graph6 Conclusion and Future WorkIn this work, we framed the inference of Metal AM process states as a next image prediction problem.

We ﬁrst trained a CNN based autoencoder to carry out the identity function of reconstruc- tion of a given image through a series of convolutions and deconvolutions. Then, we augmented the obtained latent vectors with process parameters to form our input at each timestep to be fed into an LSTM model for time series prediction. We used these two models to carry out the next image prediction in two stages: latent vector prediction with the LSTM model and image reconstruction from the decoder of the autoencoder model. Lastly, we simulated anomalies resulting from holes and our model was able to detect these anomalies through spikes in the response obtained from the selected metrics of mean squared error and structural similarity.The next step in this work is to apply the Encoder/Decoder and LSTM models to physical AM systems.

The intent of using the simulation images rather than raw simulation data for the original model training is to facilitate the transition to real in-process camera data. Two promising applications of the present work are:Anomaly detection: Using methods analogous to the one presented in Section 3 on real in-processimagery, such as a welding camera, will enable real-time anomaly detection.Internal state inference: While the present work focuses on surface-facing data, as captured by a simulated camera, the simulation contains data throughout the volume. Completing the link between the observed surface states and the behavior of what is happening below the surface will allow for exploration of the root causes of the anomaly detection. Understanding the causes of a problem create the opportunity to develop in-situ corrective strategies.1306References[1] Benjamin A Kwapong, Richard Anarﬁ, and Kenneth K Fletcher.

E INFLUENCE OF A DYNAMICALLY OPTIMIZED GALVANO BASED LASER SCANNER ON THE TOTAL SCAN TIME OF SLM PARTSS. Bulsa, T. Craeghsb, S. Clijstersa, K. Kempena, J. Sweversa, J.-P. Krutha,aUniversity of Leuven (KU Leuven), Department of Mechanical Engineering, Celestijnenlaan 300B, 3001 Heverlee, Belgium bMaterialise, Technologielaan 15, 3001 Leuven, BelgiumAbstractMost commercially available Selective Laser Melting (SLM) machines use galvano based laser scanner deflection systems. This paper describes the influence of the dynamical optimization of such galvano based laser scanner on the total scan time. The system identification of a galvano laser scanner was performed in combination with the development and implementation of an optimal ‘Input Shaper’. Tests were performed on lattice structured SLM parts. The process time was hereby compared, with and without the use of the optimal ‘Input Shaper’.

Tests were performed on lattice structured SLM parts. The process time was hereby compared, with and without the use of the optimal ‘Input Shaper’. Significant scan time reduction was observed when using the optimal ‘Input Shaper’.IntroductionSelective Laser Melting (SLM) is an Additive Manufacturing technique which enables the production of complex functional metallic parts with good mechanical properties. A schematic set-up of a typical SLM machine is shown Figure 1. In the SLM process, first, a thin layer of metal powder is deposited on a build platform by means of a powder coating system. After depositing, the powder layer is melted selectively according to a predefined scanning pattern, by a laser source [1] and a laser deflection system. After scanning a layer, the build platform moves down over a fixed distance equal to the thickness of one powder layer (in SLM typically 20 to 40 µm) and a new layer is deposited and scanned.

After scanning a layer, the build platform moves down over a fixed distance equal to the thickness of one powder layer (in SLM typically 20 to 40 µm) and a new layer is deposited and scanned. The sequence of depositing and scanning is repeated until the part(s) is (are) fully built.Figure 1: Schematic overview of SLM processIn recent years, the SLM technology has made an enormous progress in machine construction, production speed and part quality. Since material properties of SLM parts are nowadays comparable to the properties of the corresponding bulk material [2, 3], applications of the process can be found in domains like the medical sector [4], in tool-making industries [5–8], machine construction, aerospace, etc.

However, for an even larger breakthrough of SLM in industries, which demands high quality parts and low production time and thus lower cost, the SLM process must be further time optimized.Most commercial SLM machines available today are equipped with a galvano based laser deflection systems (galvano scanner) to deflect the laser according to the predefined scanning pattern. Throughout this paper it will be shown that the dynamical optimization of the galvano scanner plays a vital role in the reduction of the total process time.Experimental SetupThe dynamical optimization was performed on an in-house developed SLM machine (LM-Q). This machine is fully controlled by the industrial NI-PXIe-1082 system in combination with the NI-PXI-7853R FPGA card from National Instruments.

This machine is fully controlled by the industrial NI-PXIe-1082 system in combination with the NI-PXI-7853R FPGA card from National Instruments. The LM-Q machine is equipped with a 300W-1064nm fiber laser and a ScanLab HurryScan25 galvano scanner.National Instruments LabVIEW in combination with the LM-Q PXI controller was used for the dynamical measurements on the ScanLab HurryScan25 galvano scanner.Mathworks Matlab was used for the system identification of the galvano scanner and the development of the optimal ‘Input Shaper’.Results and discussionWhen scanning a SLM part, the machine controller sends a series of vectors to the galvano scanner. In essence only two types of vectors are used, namely ‘jump-vectors’ and ‘scan-vectors’. When executing a jump- vector, the mirrors will move as fast as possible to the new coordinate without laser power, whereas when executing a scan-vector, the mirrors will move to the new coordinate with a predefined scan speed and laser power.

The desired scan pattern is determined by the combination of jump- and scan-vectors.Modern galvano scanners are designed with an optimized PID control loop to guarantee correct movement of the mirrors. This PID control loop is specifically designed for the physical properties of the galvano motors and mirrors (damping, inertia,…). However, even with the implementation of such a PID control loop, the position of the mirrors will suffer from oscillations due to the high dynamical excitation of a (fast) jump-vector. Figure 2 shows the dynamic effect of a jump-vector. The real position follows the desired position after a specific (acceleration) time. At the end of the jump-vector, the oscillations are clearly visible. To compensate for this unwanted effect a fixed delay (jump-delay) must be used (typically 750µs to 1ms) until the oscillations die out. Executing a new vector is only possible after this ‘dead time’.

To compensate for this unwanted effect a fixed delay (jump-delay) must be used (typically 750µs to 1ms) until the oscillations die out. Executing a new vector is only possible after this ‘dead time’. This jump-delay results in huge time loss, mainly when manufacturing lattice structures containing a lot of very short scan vectors and where the total scan time is dominated by jump-vectors/jump-delays.Figure 2: Jump-vector response [9]A better solution is the use of a ‘preshaping’ method where all vectors are transformed in such a way that the system vibrations are kept to a minimum [10]. This can be achieved by the development of an ‘Optimal Input Shaper’. When implemented, the Input Shaper will transform all the vectors before sending them to the galvano scanner. Figure 3 shows a schematic overview of the control setup. The development of the ‘Input Shaper’ on the LM-Q SLM machine will be further explained in the next four sections.Figure 3: Schematic control overview1.

The development of the ‘Input Shaper’ on the LM-Q SLM machine will be further explained in the next four sections.Figure 3: Schematic control overview1. System IdentificationThe dynamic behavior of the galvano scanner was obtained by exciting and recording of the actual position of one of the galvano motors with a chirp signal. The excitation chirp signal has a start frequency of 10Hz and a stop frequency of 2kHz. The amplitude was set to 100bit with a measurement period of 1s. The system model of the galvano motor was estimated with the response behavior on the chirp signal and a ‘non- linear least squares’ system estimator.As a result of the ‘non-linear least squares’ system estimation, the following model was obtained:Figure 4 shows the Bode plot of the measured system and the estimated model.Figure 4: Bode plot of estimated galvanoscanner system2. Optimal ‘Input Shaper’The system model of the galvano motor is further used in the development of the Input Shaper.

Optimal ‘Input Shaper’The system model of the galvano motor is further used in the development of the Input Shaper. The Input Shaper is developed by inverting the system model of the galvano motor. In addition, four extra poles are placed at 3000rad/s (6 times the cutoff frequency) for system stability. After this transformation the following Input Shaper was obtained:Figure 5 shows the Bode plot of the galvano motor system model, the inverted galvano motor system model and the obtained Input Shaper with four added poles for system stability.Figure 5: Bode plot of estimated galvanoscanner system, inverse of galvanoscanner system, designed input shaper3. ValidationAfter the implementation of the Input Shaper, the system was validated with and without activating the Input Shaper. Figure 6 shows the response behavior (starred green line) of the galvano motor on a jump- vector (green line).

Figure 6 shows the response behavior (starred green line) of the galvano motor on a jump- vector (green line). When the Input Shaper is activated the jump-vector is transformed by the Input Shaper and becomes a transformed jump-vector (blue line). The response behavior on the transformed jump-vector is represented by the starred blue line. When comparing the two response behaviors, it can be noticed that both the position lag and the settling time have reduced by a factor two. Due to the reduction in position lag the jump delay can also be reduced by factor two.Figure 6: Position validation of the Input ShaperFigure 7 portrays the representation of the speed response of the galvano motor with and without theactivation of the Input Shaper. The blue line represents the velocity profile when the ‘Input Shaper’ is activated; the red line represents the velocity profile when the ‘Input Shaper’ is not activated.

The blue line represents the velocity profile when the ‘Input Shaper’ is activated; the red line represents the velocity profile when the ‘Input Shaper’ is not activated. When the Input Shaper is activated, the nominal speed is achieved 25% faster.Figure 7: Speed validation of the Input Shaper4. Test caseAs test case, a lattice structured ‘Stanford Bunny’ [11] with open pores, a strut thickness of 250µm and a volume fraction of 5% was produced: see Figure 8. Two jobs were executed, one with and one without activation of the ‘Input Shaper’. The jump-delay was reduced from 750µs to 250µs. During these jobs, the actual position of the galvano motor was logged and the total scan time was calculated.

The jump-delay was reduced from 750µs to 250µs. During these jobs, the actual position of the galvano motor was logged and the total scan time was calculated. The implementation of the ‘Input Shaper’ results in a total scan time reduction of 10% for this lattice structured bunny, from 4686 to 4219 sec.Figure 8: Test case - Porous ‘Stanford Bunny’ [scale 1:1]ConclusionThe galvano scanner of KU Leuven’s LM-Q SLM machine was dynamically optimized. This was done by the development and implementation of an optimal ‘Input Shaper’ based on the system identification of the galvano motors. Due to the higher dynamics, nominal speed can be achieved twice as fast; furthermore the ‘position lag’ is reduced by 50%. The greatest benefit of the optimal ‘Input Shaper’ is noted when manufacturing lattice structured parts. In the test case of the ‘Stanford Bunny’ total scanning time was reduced by 10% from 4686 sec to 4219 sec.References[1] J.-P. Kruth, G. Levy, F. Klocke, and T.H.C. Childs.

In this study, the influence of two build parameters (build orientation and hatch angle) on the compressive properties of 304L stainless steel was evaluated. SLM 304L samples were manufactured using three hatch angles, 0°, 67°,105° and two orientations, z-direction and x-direction, and tested using a compression frame according to ASTM E9-09. Bulk density was measured according to ASTM C373-17 before compression. Properties evaluated were the bulk density, yield strength, strength at 15% plastic-strain and strength at 30% plastic-strain. Results showed that bulk density varied minutely with respect to variation in hatch angle and build orientation, but compressive yield strength and plastic flow stress were strongly influenced by these two process parameters.

Highest compressive yield strength was measured when samples were built in the x-direction using hatch angle 67°.Keywords: Additive manufacturing; SLM; stainless steel; hatch angle; build orientationIntroduction Selective Laser Melting (SLM) is a type of additive manufacturing method that has gained greater interests over the last years due to its cost, flexibility and build capacity (i.e. ability to manufacture parts with extremely complex geometries). However, across board implementation of this manufacturing technique is lacking due to the inability to accurately predict the mechanical properties of the produced part. Process parameters such as; material parameters, build parameters etc., are known to greatly influence the mechanical properties of SLM parts. Understanding the interactions between these parameters as well as their effects on microstructure and mechanical properties is very important to researchers.

Understanding the interactions between these parameters as well as their effects on microstructure and mechanical properties is very important to researchers. Several investigations have been carried out with the aim of understanding the influence of some of these parameters on mechanical response[1]–[3].Guan et al [1] studied the effects of the build parameters, hatch angle, build direction, slice thickness and overlap rate, on the tensile properties of SLM stainless steel. They observed that1327tensile properties of SLM parts are dependent on the interval number ‘N’ and the build orientation, and samples built vertically with a higher number of deposited layers exhibited the best tensile properties. Popovich et al [2] investigated the causes of anisotropy in the SLM process.

Popovich et al [2] investigated the causes of anisotropy in the SLM process. It was concluded that process parameters played an important role in anisotropy because they controlled arrangement of planes of pores, residual stresses and grain growth (elongation and direction) in parts produced. Tolosa et al [3] investigated the effects of build orientation on the tensile properties of SLM 316L stainless steel. It was concluded that regardless of the build orientation the tensile properties of the SLM stainless steel was always exceeded that of 316L stainless steel.304L stainless steel is a type of austenitic stainless steels which exhibits good corrosion resistance as well as excellent mechanical properties. Mechanical properties of 304L stainless steel are dependent on the processing conditions as well as the loading type [5]. SLM a layer-by-layer selective melting process has a series of process parameters which may affect the final 304L stainless steel mechanical properties.

SLM a layer-by-layer selective melting process has a series of process parameters which may affect the final 304L stainless steel mechanical properties. A lot of work has been published with major focus on understanding the effects of these process parameters on tensile properties of 304L stainless steel [5]–[7] but little work has been done with respect to compressive properties hence the purpose of this study. In this study 304L stainless steel was quasi-statically compressed to study the effects of hatch angle, 0°, 67° and 105°, and build orientation, horizontal and vertical, on the compressive response.Material and Experimental ProcedureMaterial and manufacturing processSelective laser melted 304L stainless steel was analyzed quasi-statically in compression. Renishaw AM 250 machine at Missouri S&T was used to fabricate octagonal parts using 304L stainless steel powder.

Renishaw AM 250 machine at Missouri S&T was used to fabricate octagonal parts using 304L stainless steel powder. The steel powder particles used ranged from 15µm - 63 µm in size and the chemical composition is shown in Table 1. The processes for fabrication using this AM machine include: (1) deposition of thin layer of powder by the recoater; (2) melting and bonding of powder layer by laser beam; (3) lowering of build platform, deposition of another layer of powder and re- introduction of laser beam. Octagons were produced by the repetition of the above-mentioned process in an Argon filled environment.Table 1:-304L Stainless steel powder chemical compositionElement Wt%Cr 18.5Ni Mn 1.4 9.9Si 0.63Cu N 0.10.09O 0.02C 0.015P 0.012S 0.004Fe BalanceThree hatch angles 0°, 67° and 105° (labeled as A, B, and C respectively) and two build orientations (x-orientation and z-orientation) were studied.

Figure 1 shows the two orientations of specimens that were manufactured. The effects of these variations were evaluated based on changes in bulk density, yield strength, strength at 15% strain and strength at 30% strain.1328Figure 1:- Build directions used to manufacture couponsBulk densityMechanical properties of materials are negatively influenced by the presence of pores. In the SLM process presence of unmelted powder, lack of fusion as well as gas entrapments are some of the major causes of porosity [8]. Density was determined using Archimedes’ Method. The compactness (density) percentage increases as the percentage of pores reduce.

Density was determined using Archimedes’ Method. The compactness (density) percentage increases as the percentage of pores reduce. The bulk density can be calculated using equation 1.∗𝐷𝐷 𝐷𝐷 − 𝑆𝑆𝜌𝜌=(1)where ρ* is the measured bulk density, D is the dry weight of the specimen, S is the saturated weightQuasi-Static Compression TestTest coupons with aspect ratio 3:1 (elastic region) and 1:1 (plastic region) were later machined from the manufactured octagons using a CNC lathe.Cylindrical coupons were prepared for compression testing using CNC lathe according to the ASTM E9-09 standard. Coupon dimensions used during this study had an average length of 26.29mm and a diameter of 8.47mm (aspect ratio of length to diameter ~ 3:1) for the large samples, while the smaller samples averaged a length of 8.79mm and diameter of 8.44mm (aspect ratio ~ 1:1). Compression tests were performed in a 370 MTS frame according to the ASTM E9-09 standard.

Compression tests were performed in a 370 MTS frame according to the ASTM E9-09 standard. Crosshead speed of the frame was fixed to obtain an initial strain rate of 5 × 10-4 s-1 in the sample. Force and displacement were tracked during tests and used to calculate and plot the stress-strain curve. The yield strength, strength at 15% plastic-strain and strength at 30% plastic- strain were extrapolated from the resulting stress-strain curves.Results and DiscussionBulk densityThe bulk densities of the investigated build variations are presented in table 2; where A1 and A2 are specimens built with hatch angle 0° in the z-direction and x-direction respectively, B1 and B21329are specimens built with hatch angle 67° in the z-direction and x-direction respectively and C1 and C2 are specimens built with hatch angle 105° in the z-direction and x-direction respectively. It can be seen that percentage porosity is maximum when the specimen is built in the z-direction and hatch angle 0°.

It can be seen that percentage porosity is maximum when the specimen is built in the z-direction and hatch angle 0°. However, it is important to note irrespective of the selected hatch angle and build orientation, the manufactured specimens were always between 98.62% and 98.81% dense which is typical for SLM parts.

However, it is important to note irrespective of the selected hatch angle and build orientation, the manufactured specimens were always between 98.62% and 98.81% dense which is typical for SLM parts. There were no outliers in the data indicating that altering hatch angle and build orientation caused no major increase in the porosity of the samples.Table 2:- Variation in bulk density between different build orientations and hatch anglesSampleHatch angleBuild orientationDensity (g/cc)A10°z-direction7.890±0.003B167°z-direction7.905±0.007C1105°z-direction7.903±0.008A20°x-direction7.904±0.006B267°x-direction7.891±0.006C2105°x-direction7.903±0.003Build orientationThe build orientation selected during the manufacturing of SLM parts plays an important role in the physical and mechanical characteristics of the final product.

By selecting the right orientation the required support volume, build time and cost is reduced while the part quality and mechanical properties increases. During this study, two build orientations were considered (Figure 1) for different hatch angles and their effectiveness evaluated based on the yield strength. Results show that the samples built in the x-direction (horizontally) had higher yield strengths when compared to the z-direction (vertical). We assume that the difference in compressive strengths (although little) occurs either as a result the direction of the load with respect to the columnar grains or plane of pores in the layers. However, microstructural analysis was not carried out on these samples therefore it is not certain. Figure 2a shows an example of a stress-strain curve for sample B1 which was measured using long coupons. Buckling was noticed which was the cause of the dropping stress not the ultimate stress of the material.

Buckling was noticed which was the cause of the dropping stress not the ultimate stress of the material. Figure 2b shows the average yield strengths for the different build orientations and hatch angles, using the same nomenclature described for Table 2. Long coupons were used (3:1 aspect ratio) for this analysis because they defined better the elastic region of the material.1330Figure 2:- Compression results from the long coupons (3:1): (a) Example of stress-strain curve for sample B1, and (b) summary of compressive yield strength of SLM 304L stainless steelHatch angleHatch angle θ° can be defined as the angle between the scanning directions of two immediate scan layers.

The interval number ‘N’ is dependent on the hatch angle and can be calculated using equation 2, where HCF is the highest common factor between θ and 360 and θ is the hatch angle.360° HCF( θ°, 360°)𝑁𝑁 =(2)Guan et al [1] observed that by increasing the value of ‘N’ the tensile properties of 304 stainless steel increased. In this study samples were manufactured with hatch angles 0°, 67° and 105° thus resulting in various values of ‘N’. Table 3 shows the number of layers after which the melting line will coincide with the direction of the first layer when hatch angles 0°, 67° and 105° are selected.Table 3:- Relationship between hatch (θ°) angle and interval number (N)Hatch angle ()067105Interval number (N)𝛉𝛉°1 °360 °24 °The relationship between ‘N’ and plastic flow stress are shown in figure 3. True stress at 15% and 30% plastic strains were compared using short samples built in the z-direction.

True stress at 15% and 30% plastic strains were compared using short samples built in the z-direction. It was deduced from figure 2 that since samples built in the z-direction exhibited lower yield strengths (when compared to x-direction) they would deform more plastically before reaching the machine limit (100 kN). Therefore, samples built in the z-direction were selected for further investigation.1331Figure 3:-Compression results from short SLM samples for different hatch angles (0º, 67º and 105º): (a) True stress-true strain curves and (b) comparison of true stress at 15% and 30% plastic strains.From the data in Figure 3, it can be seen that Samples A showed the lowest flow stresses when compared to other configurations. Additionally, the deviation in the flow stress for A is higher when comparing the repetitions. This could show problems of the manufacturing process using this configuration. For Samples B and C, flow stresses are similar, especially at low plastic strains.

This could show problems of the manufacturing process using this configuration. For Samples B and C, flow stresses are similar, especially at low plastic strains. In average, flow stresses are only 3% higher in B than C. However, deviations in C are higher than B, especially at higher plastic strains.Conclusion In order to investigate the effects of build parameters on the compressive properties of SLM 304L stainless steel, samples were manufactured with three hatch angles (0°, 67°,105°) and two orientations (z-direction and x-direction) using the Renishaw AM 250 machine at Missouri S&T. The main summary from this study are as follows:There exists a relationship between the bulk density and the investigated build parameters.

The main summary from this study are as follows:There exists a relationship between the bulk density and the investigated build parameters. However, due to the minute variations recorded during this study it was concluded that the hatch angle and build orientation are not major controllers of bulk density in the SLM of 304L stainless steel.Hatch angle and build orientation strongly influences the compressive properties of SLM 304L stainless steel. During this study it was observed that the highest yield strength and flow stresses were recorded when hatch angle 67° was selected. This was attributed to the reduction in single direction grain growth and defect generation in the samples caused by the high interval number ‘N’.

This was attributed to the reduction in single direction grain growth and defect generation in the samples caused by the high interval number ‘N’. Samples built horizontally (x- direction) also showed higher yield strengths when compared those built vertically (z-direction).1332Acknowledgments Support from Intelligent Systems Center (ISC) and Center for Aerospace Manufacturing Technologies (CAMT) at Missouri University of Science and Technology is gratefully acknowledged. The authors would also like to acknowledge Austin Sutton for his contributions, and Honeywell Federal Manufacturing and Technology for use of the Renishaw machine.References[1] K. Guan, Z. Wang, M. Gao, X. Li, and X. Zeng, “Effects of processing parameters ontensile properties of selective laser melted 304 stainless steel,” Mater. Des., vol. 50, pp. 581–586, 2013.[2] A.A. Popovich, V. S. Sufiiarov, E. V. Borisov, I. A. Polozov, D. V. Masaylo, and A.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceTHE INFLUENCE OF CONTOUR SCANNING PARAMETERS AND STRATEGY ON SELECTIVE LASER SINTERING PA613 BUILD PART PROPERTIESC. Kummert*, H.-J. Schmid*Direct Manufacturing Research Center (DMRC) and *Particle Technology Group University of Paderborn, GermanyAbstractQualification of new materials for the laser sintering (SLS) process includes development of suitable parameters in terms of optimal part properties. Especially laser scanning parameters influence part porosity and therewith mechanical performance. In the present work tensile specimens were built of PA613 a new LS polyamide delivered by Evonik and was processed on an EOSINT P396. As build parameters have to be developed for the new material, scanning parameters and strategy of the PA613 specimen contour were varied in different ways.

As build parameters have to be developed for the new material, scanning parameters and strategy of the PA613 specimen contour were varied in different ways. Resulting part properties were investigated by XCT-analysis as well as by tensile tests. The three-dimensional part porosity, pore density and arrangement were analyzed in relation to used laser scanning parameters and resulting mechanical properties. The investigations help to understand the existing correlations between laser energy input, part porosity and mechanical performance and therewith to find optimized build parameters for the new material.IntroductionThe material portfolio for the Laser Sintering Process (SLS) is still limited and therefore the potential to find new application fields for the Additive Manufacturing process is challenging. In particular, industries like electronics, aircraft and automotive demand high performance plastics showing high strength and temperature resistance.

In particular, industries like electronics, aircraft and automotive demand high performance plastics showing high strength and temperature resistance. In this work, a new SLS material PA613, developed by Evonik, was investigated. This laser sintering powder may be implemented where PA6 is used in conventional manufacturing processes. Due to a higher melting temperature in comparison to standard laser sintering material PA12, PA613 may be applied in high temperature applications. Despite the higher melting temperature, PA613 is still processable robustly on an EOS P396 laser sintering system. However, process parameters in order to achieve high part quality have to be examined. X-ray Computer tomography (XCT) is a suitable non-destructive method to examine the build part microstructure, since high porosity and ragged surface appearance cause bad mechanical properties [1].

X-ray Computer tomography (XCT) is a suitable non-destructive method to examine the build part microstructure, since high porosity and ragged surface appearance cause bad mechanical properties [1]. Therefore, a thorough investigation of these part properties in dependence on laser scanning parameters and strategy will allow for achieving improved part quality by application of optimal laser parameters. Since energy input into the part contour is supposed to be particularly important for the specimen quality, this study was mainly dealing with contour writing.The porosity of build parts is defined as the ratio of volume of pores and the total build part volume [2]. Air pockets are also known from other polymer manufacturing processes like injection moulding, e.g. due to insufficient deaeration and schrinkage. The laser sintering process entails1582several possibilities of pore formation.

due to insufficient deaeration and schrinkage. The laser sintering process entails1582several possibilities of pore formation. Besides incomplete coalescence of polymer particles, release of gases due to polymer degradation may be an additional reason for pore formation. Following, the avoidance of porosity requires a balanced energy input by the process temperature and laser, in correlation to the particular polymer and powder bulk. To find suitable laser sintering parameters for the PA613 material with particular consideration of the influence of the microstructure, parameters for the energy input into the build part contour are varied and correlated to porosity as well as mechanical properties.State of the ArtIn the laser sintering process different material properties, parameters and therewith physical and chemical phenomena influence the resulting build part properties.

Every process step from powder layer recoating, heating up of the powder surface, laser exposure to solidification of the sinking layer is essential to generate a high density part with high strength and smooth surfaces. X- Ray Computed Tomography (XCT) has lately become an appropriate method to investigate porosity and the development of pores.Dupin et al [3] investigated the microstructure of two PA12 build parts which exhibit different particle size distributions. The authors observed 2D sections and stated different porosities in X- ray tomography for different particle size distributions, morphologies and therewith lower powder bed density which led to different welding between two successive layers where pores were localized. It is explained by a hindered laser beam penetration. Further they mention bad coalescence processes of the polymer particles and diffusion of gases into the molten polymer during cooling stage as reasons for pore development.

Further they mention bad coalescence processes of the polymer particles and diffusion of gases into the molten polymer during cooling stage as reasons for pore development. As a conclusion they highlighted particle size distribution, particle morphology, laser energy density and the crystallization temperature as the main influencing factors for residual porosity. On top of that crystallinity is investigated as another impact on mechanical properties.Also Dewulf et al. [4] investigated PA12 build part porosity in dependence of laser scanning parameters with the help of X-ray computed tomography. Beside porosity and pores volume distribution, the XCT-model was sliced and attributed to the original STL-file to calculate an average porosity per slice. Building a cube sample they varied the laser parameters laser power, hatch distance and scanning speed and therewith the energy density in four steps.

Building a cube sample they varied the laser parameters laser power, hatch distance and scanning speed and therewith the energy density in four steps. They correlated a mid-high laser energy density (Andrew Number) to minimized part porosity. Density increases with higher laser power because of higher fusion of polymer particles. Too high energy density leads to degradation of the material which causes gas emission. Furthermore, smaller pores show coalescence forming larger pores. A decrease of hatch distance leads to reduced porosity and magnitude of the porosity fluctuations. Furthermore, they investigated cylindrical pins built with different scanning patterns in different sizes, partly including contour and hatching scan lines. Exposure of only one spot causes voids in the center of the specimen. However, pins with contour scanning showed a ring of pores right inside the dense border.

Exposure of only one spot causes voids in the center of the specimen. However, pins with contour scanning showed a ring of pores right inside the dense border. Likewise, the scanning order and beam offset of the contour influences the porosity of the pins.In Rüsenbergs [5] work, thin sections were analyzed by XCT-scans and showed a denser outer area in comparison to the inner build part. This effect was also observed by Rouholamin and Hopkinson [6] who varied the laser power in in order to examine its effect on build part density, morphology and mechanical properties as well. The presence of denser structures at the edges1583compared to the centers may result in improved mechanical properties, as the averaged pore diameter was not affected by increasing laser power.Pavan et al. [7] investigated the dimensional accuracy of laser sintered PA12 parts with XCT varying the beam offset and laser power of the contour vector.

[7] investigated the dimensional accuracy of laser sintered PA12 parts with XCT varying the beam offset and laser power of the contour vector. The 3D voxel model was compared with the original CAD file and dimensional deviations were analyzed for holes in vertical and horizontal walls with different thicknesses. While the number of slices affects the dimensions of holes along the z- build direction, the xy- direction dimensions are affected more by the beam offset. It was found to be difficult to find one beam offset, which results in similar dimensions in all orientations and placements of holes within a test specimen.Ghita et al. [8] used Micro-Computer Tomography (µCT) to investigate tensile specimens of virgin and mixed PEK powder . The usage of recycling material led to higher porosity in the part and therewith to worse mechanical properties.

The usage of recycling material led to higher porosity in the part and therewith to worse mechanical properties. Coalescence of polymer particles is hindered by increased particle sizes and higher melt viscosity of the aged material, with the result of higher porosity. The latter might be compensated by higher laser energy input.On one hand material properties like particle size distribution, particle morphology and ageing stage have an impact on pore development during laser sintering and on the other hand process parameters like laser energy input and temperature are influencing factors. However, the found correlations are not linear and suitable process parameters have to be found for each new material with individual polymer and powder / particle properties.Experimental methodsWithin the present work PA613, a new polyamide for the laser sintering process, delivered by Evonik is investigated.

Virgin powder is processed on the laser sintering system EOS P 396 to build zx-direction tensile specimens according to DIN EN ISO 527 (1A). Within preliminary examinations of laser scan parameters of the filling lines the material showed quite constant mechanical properties over a wide range of energy input (0.027-0.045 J/mm² areal energy density). Where the areal energy density (ED) also called Andrew Number is calculated by the following equation:𝑃𝑃 𝑣𝑣 ∗ 𝑑𝑑 Where P is the laser power, v the scan speed and d the hatch distance between two laser scan lines. To improve the anisotropy between xy- and zx- build directions as well as the build part surface, the laser parameters and strategy of the contour is varied, whereas the filling is kept constant in the present work: Using a constant layer thickness of 120 µm the laser power to exposure the filling lines was 30 W, the scan speed 4750 mm/s and the hatch distance 0.16 mm, resulting in an area energy density (ED) of 0.0395 J/mm².

Furthermore, the contour line, shown in red in Figure 1, was varied as outlined in Table 1. The first contour parameter set “REF” serves as a reference. In comparison to that, energy density was increased by a reduced scan speed (“SS1”), higher laser power (“LP”) or a double exposure (“DE”). For parameter set “BO” the beam offset𝐸𝐸𝐸𝐸 =1584was increased and therewith energy density decreased. On top of that, further specimens were built with no contour exposure (“NC”) at all.

On top of that, further specimens were built with no contour exposure (“NC”) at all. To calculate the energy density of the contour exposure the hatch distance is replaced by the beam-offset.Figure 1: Laser scanning lines of the build part filling (green) and contour (red)Table 1: Laser parameters for contour line variationsParameter setLaser power [W]Scan speed [mm/s]Frequency [-]Beam- Offset [mm]Energy density [J/mm²]REF (reference) SS1 (decreased scan speed) SS2 (destructed by test)2525253000100010001110.20.20.20.0420.1250.125LP (increased laser power DE (Double exposure) BO (increased Beam-Offset) NC contour)(no40252530003000300012100.20.20.350.0672 x 0.0420.024For each parameter set five tensile specimens were built, whereof four were tested according to DIN EN ISO 527 on an INSTRON 5569 EH universal testing machine. The fifth of each set was used for XCT-Scans.

The fifth of each set was used for XCT-Scans. To analyze the effect of tensile destruction on the microstructure also a tested specimen built with parameter set SS1 was scanned and is called SS2 in the following. The XCT-scan (computed tomography) was performed with GE Phoenix Nanotom S with 70 kV1585150 µA and 1440 projections. As it is not possible to measure the whole tensile specimen, a cut was taken from the samples as shown in Figure 2. Furthermore, the data are analyzed using the “Defect Analysis” of the software VGstudio MAX 2.2 from Volume Graphics GmbH. The algorithm compares the gray scale value with a threshold, so that only defined defect sizes are taken into account. Hereby the voxel size is 10 µm. Furthermore, the Region-Of-Interest (ROI), which is analyzed in detail, was restricted to the contour area. It has to be noted that this region might differ between the specimens as dimensions are different due to varied contour energy input.

It has to be noted that this region might differ between the specimens as dimensions are different due to varied contour energy input. However, evaluated characteristics like porosity, averaged pore volume and pore density are relative values so the information is still representative.Figure 2: Cut-out of tensile specimen (DIN EN ISO 527 1A) for XCT-scansResults and discussionTensile specimens were built with different contour laser parameters as described before and tensile tests as well as XCT-scans were conducted. If not mentioned differently, evaluated results refer to the contour regions as the porosity of the inner specimen body was in average 4.39± 0.26 % for the seven investigated specimens. However, the contour region showed lower porosity in comparison to the inner build part for all investigated specimens as visible in Table 2. This effect was already stated by Rouholamin et al. [6] and Rüsenberg et al.

This effect was already stated by Rouholamin et al. [6] and Rüsenberg et al. [5] for PA12, who assumed that the thickness of the laser sintered part and the cooling rate might have a strong influence on the outer solid area and the porosity.1586Table 2: Result of defect analysis of contour regionsSampleREFPorosity [%]1.86Averaged pore volume [µm³] 51408Pore number concentration [1/mm³] 362SS13.08119458261SS23.58118015300LP0.8442155199DE1,6152978304BO2,0956872368NC2,1772351300Figure 3 shows a section of a XCT-scan of specimen “REF”, where the contour region is marked by the blue frame. In the contour area the pores have been evaluated (yellow). In comparison to the inner part (black pores) the number of pores seems to be decreased. The evaluated values confirm the visual appearance.

In comparison to the inner part (black pores) the number of pores seems to be decreased. The evaluated values confirm the visual appearance. The reference contour laser exposure REF shows a porosity of 1.86 % with a pore-density of 362 1/mm³ and an averaged pore volume of 51408 µm³.Figure 3: XCT-scan picture of specimen “REF”The specimen LP was built with increased contour laser power (40 W instead of 25 W) and consequently with higher energy density of 0.067 J/mm². The contour region shows the lowest porosity of all specimens due to both, lower number concentration and averaged pore volume. This effect leads to the assumption, that polymer particles show superior melt and coalescence behaviour by the higher energy input.

This effect leads to the assumption, that polymer particles show superior melt and coalescence behaviour by the higher energy input. Presumably as a consequence of the low porosity elongation at break and tensile strength are slightly improved in comparison to REF (see Table 3).1587Table 3: Tensile properties for different contour build parameter setsParameter setREFElongation at break [%] 13.59Modulus of elasticity [MPa] 2090Tensile strength [MPa]57.84SS2 (SS1)10.45213256.56LP17.77190658.78DE21.46207659.51BO14.06196858.86NC6.85186854.16However, the highest mechanical properties show specimens built with double contour exposure (DE). Even though the specimencs built with higher energy input by double contour exposure show only slightly lower porosity than the REF samples.

Even though the specimencs built with higher energy input by double contour exposure show only slightly lower porosity than the REF samples. Here another effect may play the decisive role, but has to be investigated in the future.A further reason for better mechanical performance might be a smoother part surface. InFigure 4 XCT-Scan sections of specimens LP, DE and NC are shown. No contour laser exposure NC shows worst surface appearance and worst mechanical properties as well compared to all other cases. The specimen surface is ragged caused by the alternating laser scan strategy in x- and y- build direction and lead like notches to earlier material failure.

The specimen surface is ragged caused by the alternating laser scan strategy in x- and y- build direction and lead like notches to earlier material failure. Although a comparison of the contour regions of NC to the other specimens is difficult due to different part microstructure, a qualitative comparison shows that porosity is quite high for NC due to a high average pore size volume.Figure 4: XCT-Scan sections of specimens LP, DE and NC to regard the surfacesAs the specimens SS1 and SS2 were manufactured with the same laser parameters, experimental data are examined together. Comparing the results of the “Defect Analysis” in Table 2, values are in the same range. SS2, which was destructed in a tensile test shows slightly higher porosity due to a higher number of pores. It was presumed that the tensile test leads to a stretch and an enlargement of the pores. However, the visual appearance does not substantiate this assumption, comparing in Figure 5.

It was presumed that the tensile test leads to a stretch and an enlargement of the pores. However, the visual appearance does not substantiate this assumption, comparing in Figure 5. The cross section of the specimen is smaller as it is expected for tested specimens. In comparison to the reference specimen REF both of them show a distinct increased1588porosity due to higher averaged pore volume. The number of pores is even higher for the reference. These bigger pores become clearly visible in the XCT-scan pictures where the pores are arranged like a ring which is visible in Figure 5. Possibly the local higher energy input by low scan speed leads to material degradation or at least condensation and therewith to gas inclusion. Such a pore ring was also detected by Dewulf et al. [4] for PA12 LS parts.

Such a pore ring was also detected by Dewulf et al. [4] for PA12 LS parts. Furthermore, mechanical properties are slightly lower in comparison to REF probably caused by the higher porosity within the contour region of the specimen.Figure 5: XCT-scan of SS1 and SS2 specimenThe last parameter which is varied, is the beam offset of the contour scan line. Instead of 0.2 mm the distance was set to 0.35 mm. Pores of the contour area of this specimen are bigger and result together with a higher number of pores to higher porosity in comparison to REF. Due to the higher distance of the contour to the filling scan lines, polymer particle coalescence is hindered. Again a ring of pores is visible in Figure 6 but not that obvious as seen with SS2.

Due to the higher distance of the contour to the filling scan lines, polymer particle coalescence is hindered. Again a ring of pores is visible in Figure 6 but not that obvious as seen with SS2. However, mechanical properties are not significantly affected by the higher porosity.1589Figure 6: XCT-scan of BO specimenAs a result, higher energy input into contour regions of the specimens leads until a certain degree to lower porosity in the particular area and further to higher mechanical properties and smoother surfaces. However, too high energy input causes pore development with higher average pore volume. Overall, the microstructure of the contour region appears to be an important factor for the build part performance.Conclusion and outlookIn the present work the contour region of laser sintered parts built with different exposure parameters and strategies were investigated by three dimensional X-Ray Computed Tomography (XCT).

It was shown, that a variation of energy input gets visible in the microstructure level and helps to find reasons for mechanical failures or improved properties. Apparently, the microstructure of the contour region of the specimen is crucial for mechanical performance as filling exposure parameters were kept constant, resulting in quite constant porosity in the bulk region. Not only porosity but also surfaces can be regarded by the XCT-analysis. As the notch effect is a known mechanism to cause build part failure, considering the specimen surface is an important part of the investigations. Considering the small sample size, interesting influences of different laser contour parameters became visible. While a moderate increase of energy density leads to lower porosity in the contour region and therewith to improved mechanical properties. Too high energy led to a porous ring, possibly due to material degradation and therewith to reduced mechanical properties.

Too high energy led to a porous ring, possibly due to material degradation and therewith to reduced mechanical properties. In the future, sample size and design of experiment of the conducted investigations have to be enlarged, to verify the found results and to get a better understanding about porosity development as well as build part surface appearance. This will narrow down the influences and limits of contour exposure for high build part quality.1590ReferencesReferences [1] J. P. Kruth, P. Mercelis, L. Froyen et al., “Binding Mechanisms in Selective Laser Sinteringand Selective Laser Melting,”.[2] S. Lowell and J. Shields, Powder Surface Area and Porosity, Chapman and Hall, London,New York, 1984.[3] S. Dupin, O. Lame, C. Barrès et al., “Microstructural origin of physical and mechanicalproperties of polyamide 12 processed by laser sintering,” European Polymer Journal, vol. 48, no. 9, pp. 1611–1621, 2012.[4] W. Dewulf, M. Pavan, T.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperINFLUENCE OF DROP VELOCITY AND DROPLET SPACING ON THE EQUILIBRIUM SATURATION LEVEL IN BINDER JETTINGTrenton Colton*, Joseph Liechty*, Alden McLean*, Nathan Crane*Department of Mechanical Engineering, Brigham Young University, Provo, UTAbstractUnderstanding the equilibrium saturation level is crucial to Binder Jetting (BJ). Saturation level influences dimensional accuracy, print time, green strength, and final material properties. Improved understanding of the saturation level can reduce development time for new materials and improve existing processes in BJ. Attempts have been made to predict saturation levels of parts with simple calculations from droplet primitives and capillary pressure.

Attempts have been made to predict saturation levels of parts with simple calculations from droplet primitives and capillary pressure. There is, however, limited experimental validation for these methods and they do not include the impact of drop velocity and droplet spacing. This study incorporates the influences of drop velocity and droplet spacing on the saturation level of the part. Drop primitives of varying droplet velocity and droplet spacing were compared. Results show that velocity impacts the feasible parameter space.IntroductionBinder Jetting (BJ) is a low-cost Additive Manufacturing (AM) process that uses a particle bed and inkjet technology. The powder is dispensed in thin layers and using inkjet nozzles to selectively bind powder in a chosen cross section, this is repeated layer by layer [1].

The powder is dispensed in thin layers and using inkjet nozzles to selectively bind powder in a chosen cross section, this is repeated layer by layer [1]. Similar to other powder bed AM methods such as Selective Laser Sintering (SLS), Electron Beam Melting (EBM), etc., BJ uses unbound powder from lower levels as a support structure for upper layers. BJ can utilize various materials such as ceramics, metals, and polymers [2]. For BJ to accelerate its implementation in industry; steps need to be taken to improve dimensional accuracy, improve mechanical properties, and decrease cost. These current limitations are influenced by the binder- powder interaction. Better understanding of the fundamental parameters of binder-powder interaction will decrease these limitations. Proper parameters for a new environment or new material can be identified quickly without extensive experimentation.The key to BJ is the binder-powder interaction.

Proper parameters for a new environment or new material can be identified quickly without extensive experimentation.The key to BJ is the binder-powder interaction. In BJ, 40-60 μm droplets impact powder with velocities ranging 1-10 m/s. The forces in the interaction are complex. Both capillary and inertial forces are significant and powder rearrangement is typical. In this work the influence of velocity and droplet spacing in the binder-powder interaction are investigated. Experiments have shown the interaction can impact dimensional accuracy, porosity, and strength [3].Binder saturation level measures the amount of void space in the powder filled by binder. Equilibrium saturation is the amount of binder that is stable in a powder bed. It is generally accepted that the binder saturation level in BJ should be below the equilibrium saturation level. Equilibrium saturation is a function of the binder, powder (size, shape), and pore size in the powder bed [4].

Equilibrium saturation is a function of the binder, powder (size, shape), and pore size in the powder bed [4]. While saturation is typically taken as a single parameter, there is actually a range of equilibrium values due to hysteresis between the imbibition and drainage curves [5]. Printing saturation level is the ratio of the amount of binder per volume of powder to the porosity per volume of powder.99Printhead speed has been shown to alter the saturation level and decrease dimensional accuracy in printed parts [6]. Printhead speed alters the droplet velocity and direction. Higher printhead speed increases splashing, rebounding, and cratering. The saturation level estimated from simple calculations based on binder drop primitives and capillary pressure have little experimental validation and do not include the impact of velocity [4, 7]. An alternate approach to understanding the interaction is in situ observation.

An alternate approach to understanding the interaction is in situ observation. Studies conducted using in situ high-speed imaging and x-ray imaging have observed deformation of the powder bed and primitive formation [8, 9].In this study, we characterize the influence of droplet velocity and droplet spacing on the equilibrium saturation level. Experiments performed utilizing sessile drop primitives, and printed lines characterize the dependence of equilibrium saturation levels on droplet velocity and droplet spacing.Fluid Flow in PowderMuch of the current understanding of the binder-powder interaction physics is based on the flow of fluid in soil. Bredt discussed the challenge presented by the scale between soil hydrology and BJ [5]. The structures in soil hydrology can be quite large and models are not bounded by time which exaggerate the predicted size of printed features.

The structures in soil hydrology can be quite large and models are not bounded by time which exaggerate the predicted size of printed features. Bredt’s model assumed no kinetic impact, no hysteresis in the advancing or receding of droplet contact angles, and no evaporation effect [5]. Binder will migrate through voids in powder until the pressure across the binder-air interface is in equilibrium. The curvature of the liquid interfaces in the pores creates a pressure difference across the interface. The pressure difference varies with saturation level.In BJ, one point on the surface is initially fully saturated. Unsaturated regions absorb binder as capillary pressure drives flow from saturated regions until equilibrium is reached. Due to powder packing variation and the hysteresis in the imbibition and drainage curves the saturation level will be slightly different throughout the wetted region.

Due to powder packing variation and the hysteresis in the imbibition and drainage curves the saturation level will be slightly different throughout the wetted region. Observations of the droplet impact such as cratering and particle movement contribute to undesirable part defects [8, 9]. This deformation of powder beds and impact of kinetic energy can be seen in the overall part quality and dimensional tolerance. Saturation decreased and inaccuracies increased with printhead speed [6].The prediction of saturation level in BJ is valuable in decreasing time to identify appropriate printing parameters. Miyanaji [7] proposed the first physics-based model to predict the equilibrium saturation level based on the measured capillary pressure. The model assumes a single equivalent pore size which amounts to modeling the powder bed as a series of cylindrical pores of the same size. Problems with this model occurred when successive layers were printed the lower layers became more saturated.

Problems with this model occurred when successive layers were printed the lower layers became more saturated. It was speculated to be caused by drainage from the upper layers [7]. Unfortunately, the limited experimental data set did not show consistent correlation between model predictions of equilibrium saturation and part-level measurements. Miyanaji attributed the difference to errors in powder property measurement. However, this model does not account for parameters such as the impact of droplet velocity in BJ printing or the range of pore sizes in a powder bed or the effects of binder evaporation between printed layers.100Bai (2019) characterized the binder-powder interaction by studying the primitives formed by millimeter-scale sessile drops. The study calculated the dynamic contact angle from measured imbibition times of sessile drops and attempted to correlate these measurements to saturation level [4].

The study calculated the dynamic contact angle from measured imbibition times of sessile drops and attempted to correlate these measurements to saturation level [4]. Bai noted that droplet velocity was a variable to account for in future testing. The sessile drop is limited in predicting the binder-powder interaction in BJ because it did not account for effects of kinetic energy. Predicting the interaction is valuable and can improve with a greater understanding of kinetic energy and droplet spacing in BJ.MethodsThe characterization of the binder-powder interaction was studied through two experimental geometries.

Initial binder mass and the mass of the resultant primitives were measured to calculate equilibrium saturation levels based on the binder mass ((cid:1865)(cid:3029)) and the primitive mass ((cid:1865)(cid:3043)), using:(cid:1845) (cid:3404)(cid:1865)(cid:3043) (cid:2025)(cid:3043)(cid:3029)(cid:1865)(cid:3029) (cid:2025)(cid:3029)(cid:1499) (cid:4666)(cid:883) (cid:3398) (cid:1842)(cid:3033)(cid:4667)(1)where (cid:2025)(cid:3029) is binder density and (cid:2025)(cid:3043)(cid:3029) is powder bed density, and (cid:1842)(cid:3033) is the packing fraction. The experimental geometries isolated the effects of velocity and droplet spacing. The BJ experimental apparatus was designed and built to control droplet velocity and droplet spacing independently for repeatable experiments.The experiments were conducted using ExOne stainless steel 420 powder with an average particle size of 30 μm. Powder beds were prepared using aluminum containers filled with powder and tapped.

Powder beds were prepared using aluminum containers filled with powder and tapped. The surface was then leveled using a straight edge. The tapped density was approximately 4.66 g/cm3. This equates to a packing fraction of 0.58. Density was calculated by measuring the mass required to fill a container with a known volume. Powder beds were prepared in this manner in all experiments.The two experimental procedures are outlined as (1) primitive formation from droplets released from varying heights to vary velocity. (2) Line printing with varying velocities and droplet spacing.The first experiment procedure, primitive droplet formation was measured by dropping millimeter-scaled droplets (2-14.5 microliters) into a powder bed from controlled heights to control the droplet velocity. Binder was dispensed from a 5 ml syringe through a needle.

Binder was dispensed from a 5 ml syringe through a needle. The droplet volume was measured by placing the powder bed on the scale (Ohaus AX224) to measure the mass of each droplet as it was deposited as illustrated in Figure 1. To average effects of any powder bed density variation within the bed, droplet primitive placement was varied for each droplet release height. Drops were released at heights of 0,10, and 20 cm directly over the scale allowing each drop mass to be measured accurately [10]. Larger release heights caused splashing. ExOne solvent binder was used and volume was calculated from mass measurements. The ExOne solvent binder had a measured density of 0.965 g/cm3. The powder beds were heated in an oven at 180 degrees Celsius for 2 hours immediately after. The primitives’ masses were measured to calculate the101equilibrium saturation level from Equation 1.

The powder beds were heated in an oven at 180 degrees Celsius for 2 hours immediately after. The primitives’ masses were measured to calculate the101equilibrium saturation level from Equation 1. This ratio normalized the values and allows for a comparison with the range of values.t BINDER DISPENSER WITH uADJUSTABLE HEIGHTBINDER Q DROPLETFigure 1: Apparatus design for droplet primitive experiment. The droplet was dropped from controlled heights above the scale. By controlling the height, the velocity of the droplet changed. Heights of 0, 10, and 20 cm were included in the test. Heights above 20 cm splashed.The second experimental procedure, the custom apparatus for line printing is shown in Figure 2. The apparatus consisted of MicroFab MJ-AB-01 single nozzle (40 and 60 μm orifice) print setup in unison with three motorized linear stages.

The apparatus consisted of MicroFab MJ-AB-01 single nozzle (40 and 60 μm orifice) print setup in unison with three motorized linear stages. The MicroFab single nozzle print setup utilized a MicroFab pneumatics control system, binder reservoir, tubing, and a piezoelectric nozzle. The binder reservoir is pressurized by the pneumatics console and maintained at the same height relative to the printhead to prevent pressure difference during movement of the printhead.The piezoelectric nozzle ejects droplets when an electric waveform signal is received from the drive box (Jet Drive III). Figure 3 shows the integration of the various systems. The droplet mass and initial velocity are dependent on the voltage and timing of the waveform. The general waveform used is known as a bipolar waveform [11]. Droplet velocity was changed by altering the voltage of the waveform. Droplet images were produced from IDS UI-3370CP-M-GL camera at 2.5 magnification and a strobing LED.

Droplet velocity was changed by altering the voltage of the waveform. Droplet images were produced from IDS UI-3370CP-M-GL camera at 2.5 magnification and a strobing LED. Distance measured between droplet images created by successive strobes flashes were used to calculated droplet velocity.A custom powder bed container was made to attach to two linear motor stages. ExOne solvent binder was used for the experiment. The inkjet nozzle was triggered in unison with the movement of the linear motor creating printed lines with the powder. To determine the amount of binder used, the system ran twenty lines, twenty mm in length with a small container underneath to collect binder. The mass of dispensed binder was measured. After printing, the powder bed was heated at 180 degrees Celsius for thirty minutes in an oven.

The mass of dispensed binder was measured. After printing, the powder bed was heated at 180 degrees Celsius for thirty minutes in an oven. Each individual lines’ mass was measured to calculate the effective saturation level as a function of droplet velocity and droplet spacing using Equation 1.102Binder ReservoirPiezoelectric NozzleCameraMotorized StagesFigure 2: Image of the experimental apparatus for line printing. Shown are stages for bed and printhead movement.COMPRESSED AIR□ COMPRESSED AIR/BINDER□ MOTION CONTROLPIEZO CONTROLFigure 3: A diagram of the experimental apparatus and connectivity between units. A LabVIEW VI drove the Newport linear stages in unison with triggering the JetDrive III for control of the printhead.103Results and DiscussionThe equilibrium saturation of drops from the first experimental procedure (Figure 1) is summarized in Figure 4.

millimeter-scaled droplets (2-14.5 microliters) with masses ranging 0.002-0.014 grams were drop heights were 0, 10, and 20 cm. The equilibrium saturation level ranged between 65% to 95% with significant variation in each measurement. The saturation level is relatively insensitive to drop height suggesting that increasing velocity did not change effective saturation levels for these millimeter-scale droplets. The variation of saturation levels at each height may be attributed to variation of powder bed density within the powder bed or changes in environmental conditions.

The variation of saturation levels at each height may be attributed to variation of powder bed density within the powder bed or changes in environmental conditions. However, there may be significant differences in the physics of the interaction when comparing single mm-scale droplets to lines of inkjet-printed droplets where each droplet is <100 microns in diameter.X0.9 r0.8 rC: 0.7 r 0 ~ 0.6 >- =, ro cn 0.5 >- (1) > U 0.4 r :jg W 0.3 r~X ~ ~ X~~~a0.2 >-0.1 >-0IIIIIIIIIII0246810121416182022Drop Height (cm)Figure 4: Saturation level of primitives from drops made using a syringe shown. No correlation is shown between low velocity and relatively large droplets.Results from inkjet printing of lines are shown in Figure 5. The trend of higher saturation level at lower velocities can clearly be seen.

The trend of higher saturation level at lower velocities can clearly be seen. There appears to be a limit at lower velocities approaching 80% saturation and upper velocities approaching 50% saturation—well beyond the standard deviation in the measurements. The range of tested droplet velocity covers the typical seen in BJ systems. Satellite droplets, rebounding, and splashing could be affecting these saturation measurements, but no clear evidence of these defects was observed inspecting the powder bed after printing.

Satellite droplets, rebounding, and splashing could be affecting these saturation measurements, but no clear evidence of these defects was observed inspecting the powder bed after printing. However, magnified imaging of the powder bed or high-speed imaging during printing might provide additional insight.104IIIIIII0.90.8 -c 0.7 - 0 ~ :i 0.6 - ro (/) E 0.5 - ::::,@ 0.4 - ·5 CT w 0.3 -IIII II0.20.1 ~0234567891011Approximate Velocity (mis)Figure 5: Results from lines printed with a 60-micron orifice print head at 2000 HZ in 420 stainless steel powder with one standard deviation of the average shown in error bars. Trends to higher equilibrium saturation at lower velocities is seen. Droplet spacing was uniform through the experiment. Droplet shape, velocity, and size varied with the change in waveform.

Trends to higher equilibrium saturation at lower velocities is seen. Droplet spacing was uniform through the experiment. Droplet shape, velocity, and size varied with the change in waveform. Error bars are small but show one standard deviation.The decrease in effective saturation levels with increasing velocity shows that the kinetic energy of the droplet impacts the absorption characteristics of the droplet—at least in the case of a single line on bare powder. Millimeter-scale droplets falling 20 cm will have a velocity of approximately 2 m/s if air resistance is negligible. The lower range of saturations in these mm- scale droplets is comparable to the mean value of the lines printed with 3 m/s droplets. However, the typical printing velocity is >5 m/s where the kinetic energy is observed to reduce the effective saturation levels of the binder in the powder.The impact of droplet spacing on effective saturation levels is shown in Figure 6.

Four separate data sets at different droplet velocity were recorded. A 40 μm orifice was used and droplet spacings ranging between 5-12.5 μm, droplet sizes ranging between 33-75 μm diameter. The trend of decreasing equilibrium saturation with increasing velocity generally holds with different droplet spacings though was broken with overlap at velocities 5.4 and 9.1 m/s. An increasing droplet spacing clearly decreased equilibrium saturation at all velocities.The impact of increased velocity is decreased with larger droplet overlap as seen in Figure 6. Images of printed lines shown in Figure 7 demonstrated that the line width increased but the edge roughness stayed similar. Uneven edges may be caused by the tendency for binder to fill the smallest pores first due to the larger capillary pressure in small pores. Possible errors such as satellite drops, rebounding, and splashing could also be playing roles in this situation.

Possible errors such as satellite drops, rebounding, and splashing could also be playing roles in this situation. However, there is good agreement between the two orifice sizes at same velocities.105IIIII0.90.8+ - 2.6m/s -+ - s.4mls 9.1 mis --+- 11.1 misai > ~ 0.7~C2 0.6 ~ :, ~ 0.5E ·E o.4 ~ -~0.3 UJ~f-f-~-f~f0.2~0.1f-0IIIII46810 Droplet Diameter/Droplet Spacing121416Figure 6: Results from differing droplet diameter/droplet spacing using a 40 micron orifice at 1000 HZ. One standard deviation of the average shown in error bars. The droplets were spaced between 5 and 12.5 microns apart. The ratio was used to normalize the effect of the changing volume from the different waveforms at each velocity.Figure 7: Bottom line is 4 m/s droplets at 1000 HZ, droplet spacing of 12.5 microns.

The ratio was used to normalize the effect of the changing volume from the different waveforms at each velocity.Figure 7: Bottom line is 4 m/s droplets at 1000 HZ, droplet spacing of 12.5 microns. Top line shown is 4 m/s at 1000 HZ, droplet spacing of 5 microns.106ConclusionThe binder-powder interaction is complex. Increased understanding of the interaction will lead to improved processes and faster development of new materials. This study examined impact of velocity on the equilibrium saturation level. The few models currently predicting the saturation level disregard the impact of velocity and assume quasi-static conditions. This work shows that droplet velocity does have a significant impact on the equilibrium saturation level. In the test material system, increasing droplet velocity from 3 m/s to 9 m/s reduces saturation from 73% to 48%.

In the test material system, increasing droplet velocity from 3 m/s to 9 m/s reduces saturation from 73% to 48%. As improved powder printing models are developed the impact of velocity should be included and further researched.Decreasing droplet spacing increases saturation towards the values characteristic of low velocity impacts. Similar effects may also be seen from printing adjacent lines or subsequent layers. The variation of saturation level with velocity and droplet spacing should be investigated further to assist in developing a more complete model of the binder-powder interaction. Observation of satellite drops, splashing, rebounding, and a changing waveform need to be investigated further to understand the mechanism for the changes.

Observation of satellite drops, splashing, rebounding, and a changing waveform need to be investigated further to understand the mechanism for the changes. Experimental observations of other geometries such as multilayer and multiline parts could provide greater insight to characterize the interaction.AcknowledgmentThis work was funded in part by the National Science Foundation through award CMMI1912028.References[1][2][3][4][5][6][7]M. Ziaee and N. B. Crane, "Binder Jetting: A Review of Process, Materials, and Methods," Additive Manufacturing, 2019/06/22/ 2019, doi: https://doi.org/10.1016/j.addma.2019.05.031. L. E. Murr, "Frontiers of 3D Printing/Additive Manufacturing: from Human Organs to Aircraft Fabrication†," Journal of Materials Science & Technology, vol. 32, no. 10, pp. 987-995, 2016/10/01/ 2016, doi: https://doi.org/10.1016/j.jmst.2016.08.011. Y. Bai, G. Wagner, and C. B.

However, the effect of different interfacial materials (due to the presence or absence of a shape converter) and the pause time during the build process on a part’s material properties is not well-understood. Therefore, the tensile strength of 3D-printed embedded specimens with and without shape converters and with different intervals of pause time is tested in this study. The results from this experimental analysis can be useful for the design guidelines for AM with embedded components as they provide an initial understanding of mechanical properties of these parts.1. INTRODUCTIONMoving from rapid prototyping towards direct manufacturing, additive manufacturing (AM, also known as 3D printing) has evolved as a manufacturing method over the past thirty years, offering unparalleled capability in terms of accessibility of technology, flexibility of design, and creation of end-use products [1].

As an inexpensive and relatively uncomplicated process, material extrusion or fused deposition modelling (FDM) is one of the most common and widely used types of AM. In material extrusion, a heated nozzle melts, extrudes, and deposits the fed thermoplastic filament on a build tray in a prescribed geometry. Solid layers are generated by side-by-side deposition and solidification of molten roads of thermoplastic. The build tray is lowered once a layer is completed. The next layer bonds with the previous layer as it cools down. The control of thermal environment is important for successful bonding of roads, and hence the system is often enclosed to maintain appropriate temperature below the melting point of the material being extruded [2].Contrary to traditional subtractive manufacturing, the layer-by-layer nature of AM allows for access to the entire volume of the workpiece throughout the build process.

This, in turn, gives an opportunity to embed sensors, actuators, circuits, and other functional components within a part during the manufacturing process [3,4]. Its applications have recently been recognized for design and production of optimized parts and sophisticated “smart” objects. Typically, the embedding process in AM involves 1) designing the cavity for the embedded component, 2) pausing the print when the top layer of the cavity is reached, 3) inserting the component and 4) resuming the printing process [4]. While in-situ embedding is a powerful tool for AM, there are two factors that may greatly affect the strength of these embedded parts: process interruption and material transition. Since thermal environment control is crucial for bonding [2], an interruption in the process causes the layer to cool down, and may affect the layer bonding when resumed. The material transition847occurs when the printing is resumed after embedding a foreign component.

The material transition847occurs when the printing is resumed after embedding a foreign component. The cavity is covered with another component and the resumed layer is deposited on it. While qualitative observation notes that the layer adhesion may be impacted by both of these factors, their effects on the overall strength of the part are unknown.In this research paper, the effects of the embedding process on the strength of material extruded parts were studied by preparing specimens with a Mark One 3D printer. The Mark One printer by MarkForged is a 3D printer capable of creating carbon fiber reinforced thermoplastic parts via dual-extrusion of a Nylon base material with continuous carbon fibers.

The Mark One printer by MarkForged is a 3D printer capable of creating carbon fiber reinforced thermoplastic parts via dual-extrusion of a Nylon base material with continuous carbon fibers. This combination of polymer printing with continuous carbon fibers enables users to create physical objects with high strength to weight ratios; the carbon fiber prints are claimed to be twenty times stiffer and five times stronger than standard ABS plastic [5], making it practical to produce high strength parts with AM. With the printer, users can also preselect a layer and insert a pause while preparing the 3D model for printing. The print bed reregisters with 10-micron accuracy, allowing users to pause a print, remove the bed, add components, reinsert the bed, and then continue the print with a high degree of precision.

The print bed reregisters with 10-micron accuracy, allowing users to pause a print, remove the bed, add components, reinsert the bed, and then continue the print with a high degree of precision. This focus on the potential of in-situ embedding in AM gives the Mark One an advantage over similar desktop material extrusion systems, and makes it a well-poised candidate for study of the impacts of in-situ embedding within AM.2. LITERATURE REVIEWPrevious research related to embedding with AM has been divided into three parts. Section 2.1 explores the diverse applications of the embedding process. Section 2.2 summarizes prior efforts in embedding materials using different AM process types. Section 2.3 outlines the research on mechanical properties of embedded parts.

Section 2.2 summarizes prior efforts in embedding materials using different AM process types. Section 2.3 outlines the research on mechanical properties of embedded parts. This includes considerations for designing embedded components and addresses concerns with mechanical properties of embedded parts.2.1 Multi-Functional parts via AMThe popularity of the Internet of Things [6] has spurred recent AM research in the validation of both form with functionality, which primarily involves embedding functional components in parts [7–10]. Technologies such as near field communication, real-time localization, and feedback from embedded sensors has enabled the transformation of everyday objects into smart objects. These objects are aware of their surroundings and can communicate or react accordingly[11,12]. Isanaka and co-authors summarized these roles of embedding via AM technologies to help establish a Cyber-Enabled Manufacturing (CEM) environment.

Isanaka and co-authors summarized these roles of embedding via AM technologies to help establish a Cyber-Enabled Manufacturing (CEM) environment. The CEM environment is a smart manufacturing environment, which consists of a network of embedded sensors, coupled with control systems, to effectively gather information and offer immediate response in manufacturing facilities [13]. AM makes embedding these sensors feasible and accessible. Integrating current technologies with AM allows for the design and production of sophisticated products in a CEM- like automated manner, and therefore with reduced product development cycle time [7,14].One of the major applications of multi-functionality is structural health monitoring (SHM), which is an emerging method for non-destructive product evaluation. It requires sensors to be embedded into the structure to capture, log, and analyze real time data about the health of the structure.

It requires sensors to be embedded into the structure to capture, log, and analyze real time data about the health of the structure. Benefits of this application are realized in structures like bridges, buildings, aircrafts and mega machines where catastrophic failures must be avoided [15]. Strantza and co-authors [16]848evaluated the performance of SHM systems embedded via AM, for their suitability in non- destructive testing of structures. The system was found reliable, and is expected to be used to obtain a novel approach to structural design that relies on lightweight structures. This rise in applications of in-situ embedding via AM in the fields requiring high safety and performance standards, require careful process and quality control.

This rise in applications of in-situ embedding via AM in the fields requiring high safety and performance standards, require careful process and quality control. Suitable standards for testing will ensure the acceptance of this rising technology [1].2.2 Embedding in different AM process typesLayer by layer material addition allows for in situ embedding applications to be realized in a wide variety of AM process types. A significant amount of the research on embedding via AM has been done with Stereolithography (SLA) process, a vat photo-polymerization AM technology [9]. Kataria and Rosen [17] demonstrated methods for fabricating embedded complex devices in SLA. Macdonald and co-authors [7] developed an SLA process to 3D print electronics, in order to reduce product development cycle time. SLA has also been used to fabricate 3D printed structures with embedded electronics for a space vehicle ‘CubesSat Trailblazer’, launched by NASA in November 2013.

SLA has also been used to fabricate 3D printed structures with embedded electronics for a space vehicle ‘CubesSat Trailblazer’, launched by NASA in November 2013. However, a lack of durability was identified due to the material properties of the photopolymer resin, which is the material used for fabrication in SLA process [14,18].When considering embedding within metal structures, traditional powder metallurgy processes involve isostatic pressing and diffusion bonding and are capable of reaching a temperature up to 565°C. This high temperature can prove damaging for sensitive embedded components, which has limited its usefulness in embedding applications [19]. In the area of metal powder-bed fusion AM, Rodriguez and coauthors [20] measured the surface temperature while printing through electron beam melting process by embedding a thermocouple in the build plate.

The electron beam can make the temperature of the metal surface to go up to 2269°C [21], making it impossible to insert overly delicate components. The sheet lamination process, ultrasonic additive manufacturing (UAM) is a solid state joining process in which thin metallic tapes are ultrasonically welded on top of one another and periodically machined to create a final part [22]. Hahnlel and co-authors demonstrated a successful electrical insulation of embedded material through UAM process on aluminum matrix composites. The parts fabricated through this process experience temperature as low as 25°C and therefore it is safe for embedded components like smart materials or sensors [19].

The parts fabricated through this process experience temperature as low as 25°C and therefore it is safe for embedded components like smart materials or sensors [19]. However, UAM has limited success bonding dissimilar materials, limiting its application in the aerospace, automotive, electrical and power generation industries [23].The material jetting process is an AM technique that utilizes drop-on-demand inkjet printing to selectively deposit droplets of photopolymer directly onto a build platform. Because of low range of working temperature and direct material addition layer after layer, this process has proven well-suited for component embedding. However, the print head assembly block passes over the printed part at a clearance of only 100 microns. Any protrusions from the embedded component can cause damage to the nozzle [4]. Also, there is only a limited number of photo polymeric materials available for material jetting.

Any protrusions from the embedded component can cause damage to the nozzle [4]. Also, there is only a limited number of photo polymeric materials available for material jetting. As with SLA, this type of material results in fragile parts.Polymer material extrusion is the one of the least explored but also potentially the most useful AM process types for in-situ embedding. With proper clearance design, components can be embedded without subjecting them to damaging temperature and pressure. While inherent limitations of material extrusion make it difficult to produce high tolerance or products with a fine849surface finish, this limitation can be eliminated by its ability to embed highly finished parts like bearings and shafts [24]. Sbriglia and co-authors [25] embedded sensors in a material extrusion part to analyze the operational choice for the process.

Sbriglia and co-authors [25] embedded sensors in a material extrusion part to analyze the operational choice for the process. The study concurred with the findings of Stark and co-authors about embedding sensors in FDM parts as a useful method for state-of-health monitoring in deployed systems [26]. Aguilera and co-authors also demonstrated an integrated process for embedding high performance conductors directly into the thermoplastic material extrusion substrate [27]. The wide variety of hybrid processes and thermoplastics available for material extrusion help to enable its wide application in various industries and make it a suitable candidate for in-situ embedding research.2.3 Mechanical Properties of Embedded PartsBecause inter-layer bond strength is a major factor when it comes to the mechanical properties of 3D printed part [28], in-situ embedding has the likely potential to impact a structure’s mechanical properties.

However, a review of the available literature shows that research in this area is still sparse. Klift, Koga, Todoroki, Ueda, & Hirano, performed tensile tests using Mark One 3D printer to compare strength differences in two layered (2CF) and six layered (6CF) continuous carbon fiber embedded within a Nylon base. The tests revealed that discontinuities of the fibers led to premature failure in the areas where fibers were absent and reduced the tensile strength of the composite material. A further investigation of cross sections of 2CF and 6CF specimens revealed that there was more void area in 6CF specimens. These voids had a negative impact on the elastic modulus of the composite, which calls for future work on effects of different arrangements of carbon fiber layers on the mechanical properties of these composites [29].

Outside of AM, a study on mechanical properties of casted epoxy resin with and without embedded silicon substrates verified that the ultimate tensile strength dramatically reduced in comparison to that of without embedding. Also, the points of failure for each specimen were located on the embedded section. It was concluded that the strength of the part embedded with an insert will have to be investigated if the insert is harder [30].2.4 ContextAs this review shows, existing literature has demonstrated a consistent growth in acknowledgement of applications of embedding via AM. This has led to further investigations into processes which could assist in embedding. However, to the best of the author’s knowledge, there is little-to-no research performed to understand the effects of process interruption or embedding on the properties of the printed parts.

However, to the best of the author’s knowledge, there is little-to-no research performed to understand the effects of process interruption or embedding on the properties of the printed parts. To address this critical gap, the purpose of this paper is to investigate the effects of the embedding process on the mechanical properties of material extrusion parts. To this end, the authors use tensile testing in order to determine i) how process interruption affects the strength of a printed part and ii) how the presence of an embedded artifact affects the strength of the printed part. Process interruption specimens are prepared carefully with different pause time intervals, to test if there is any correlation between print pause interval and mechanical properties, while embedding specimens include both those with shape converters and without shape converters.

Knowledge of the influence of these parameters can greatly benefit in making design decisions for manufacturing systems or for preparing safety critical parts. Section 3 of this paper discusses the experiment performed to findings on effects of process interruption on strengths of the parts. Section 4 discusses the effects of embedding process found through experimentation on the strength of the parts. Section 5 summarizes the findings of this research paper and discusses the gaps and future work in order.8503. EFFECTS OF PROCESS INTERRUPTION3.1 Experimental SetupProcess interruption during printing is a requirement while embedding, but these interruptions can also be caused by power outages, system errors or material shortages. When dealing with such issues, the decision of whether to resume the print or to restart the print, lies with the system user. To make an informed decision, it is necessary to know the impact of this interruption on the strength of the printed part.

To make an informed decision, it is necessary to know the impact of this interruption on the strength of the printed part. To better understand these effects, a Mark One 3D printer is used to create specimens from nylon filament. The system can achieve a layer resolution of 100 microns and Eiger (Mark Forged’s cloud based, print preparation software) gives a provision to insert a pause after any selected layer. This allows for a precise consistency in specimens prepared. All the specimens were printed with z axis orientation (Figure 1) so as to study the magnified effects of process interruption on tensile properties of the parts.(a)(b)Figure 1. (a) Build orientation of the specimen on Mark One and (b) Mark One while printing the specimen.Standard ASTM specimen designs were attempted for print [31], however, the specimen dimensions were either too thin for a successful print (Figure 2a), or too tall to fit within the build volume.

After several attempts at specimen design, the final design (Figure 2b and 2c) was selected based on physical limits of the build platform, success of the print, and careful consideration of parameters to be studied. The strength along z- axis is lower for material extrusion parts as compared to other build orientations. This material anisotropy is an inherent limitation of material extrusion process, as material likely cools more thoroughly between layers than between roads, potentially affecting fusion of the thermoplastic [32]. Therefore, specimens prepared in this orientation will be more sensitive to the effects of process interruption. This allows authors to study if the embedding process introduces a weakness on the layer where the pause occurs.Specimen design for different treatments (no pause and pause with time duration of 5, 15, 30 and 60 minutes) are shown in Figure 2b and 2c.

The reference specimens, with 0 minutes of pause time, were designed with the dimension of 12mm x 24mm x 96mm and were printed without any process interruption. For process interruption, specimens were printed with a pause at the center layer using Eiger (as shown in Figure 2c). These specimens were prepared with pause time intervals of 60 minutes, 30 minutes, 15 minutes and 5 minutes.851(a)(b)(c)Figure 2. Shows (a) the failed print of a standard ASTM dogbone sample design, (b) the final design of specimen for no pause and (c) with pause of different time durations.The build was prepared using the Eiger software with 0.2 mm layer height, 100% fill density, and rectangular fill pattern. Each specimen was printed individually and stored in a dark location at room temperature (24°C), with 50% humidity. Desiccant packets were used in an attempt to control the impact of humidity. The testing was performed after at least 48 hours of storage.

Desiccant packets were used in an attempt to control the impact of humidity. The testing was performed after at least 48 hours of storage. The final printed specimens are shown in Figure 3.(a)(b)Figure 3. (a) A single printed specimen for tensile testing and (b) all specimens prepared and marked for indicating treatments, before tensile testing.Tensile testing was performed using Instron 5866 Mechanical Testing Machine, using mechanical wedge action grips on a 5kN load cell (Figure 4). The pull rate was 5mm/min. Serrated jaw faced grips of thickness range 6.35mm to 12.7mm were used. Gauge length, width, and thickness of each samples were recorded in millimeters and every 0.1 second the extension (mm), tensile stress (MPa), load (N), and tensile strain (mm/mm) were recorded. The temperature during testing was recorded as 24°C and the humidity was recorded as 55%.852Figure 4.

The temperature during testing was recorded as 24°C and the humidity was recorded as 55%.852Figure 4. Tensile testing on Instron 5866 Mechanical Testing Machine.3.2 Results and discussionThe statistics of the maximum tensile stress data obtained are reported in Table 1. Threespecimens were tested for each treatment.Table 1. Statistics of maximum tensile stress (MPa) values obtained for each treatment:Treatment 0 minute pause 5 minute pause 15 minute pause 30 minute pause 60 minute pauseMean 13.03 7.43 7.00 6.27 7.54Median 12.92 7.90 7.03 6.13 7.17Std. Deviation Variance0.202 1.42 0.75 0.69 0.750.04 2.01 0.57 0.48 0.56Range Minimum Maximum 0.36 2.72 1.51 1.36 1.3512.90 5.84 6.23 5.66 7.0513.26 8.56 7.740 7.02 8.40To perform statistical analysis on the data obtained so as to find the significance in the differences recorded, the data was checked for the assumptions for parametric testing.

These were, independence of cases, normality, and homogeneity of variances. All the assumptions were met except for normality. The Shapiro-Wilk test (Table 2) confirmed that the data for Maximum Tensile Stress is not normal (p=0.002, < 0.05). Therefore, non-parametric tests were performed to analyze data for maximum tensile stress.Table 2. Tests for NormalityMaximum_Tensile_StressKolmogorov-Smirnova df 15Statistic .254Sig. .010Statistic .786Shapiro-Wilk df 15Sig. .002a. Lilliefors Significance CorrectionSince there are five groups (treatments) to be compared, a Kruskal-Wallis H test (also known as one-way ANOVA on ranks) was performed to identify if there was any statistically significant difference in the data due to treatments. A Kruskal-Wallis H test is a rank-based nonparametric test used to determine if there are statistically significant differences between two or more groups of an independent variable on a continuous or ordinal dependent variable.

The test showed that there was a statistically significant difference in Maximum Tensile stress values853between different treatments, χ2(2) = 9.633, p=0.047, with a mean rank of 14.00 for no pause, 8.00 for 5 minutes pause, 6.33 for 15 mins pause, 3.00 for 30 mins pause and 8.67 for 60 mins pause. Effect size quantifies how well this difference works in a range of contexts. This value can be calculated as shown below, where N is the total number of observations.Effect size = χ2/(N-1)=9.633/(15-1) =0.688The value 0.688 of effect size means that 68.8% of the variability in rank scores is accounted for by different treatments. The Kruskall Wallis test only verifies that at least two groups were significantly different. To identify where these differences occurred, post hoc analysis was performed by pairwise comparision of each combination of groups (based on treatment).

To identify where these differences occurred, post hoc analysis was performed by pairwise comparision of each combination of groups (based on treatment). To study the individual tensile property effects of process interruption while embedding (pausing the print to insert) or due to system errors, the specimens prepared with pause were compared by grouping them on the basis of time interval of pause. The reference values were set to be the data obtained for specimens with no process interruption (0 minutes pause). The maximum tensile stress values are reported by pause time intervals in Figure 5. Error bars indicate sample standard deviations of the data points for each treatment.14S S E R T S E L I S N E T M U M X A MI) A P M(1210864205051 01 52 03 0 PAUSE TIME IN MINUTES2 53 54 04 55 05 56 06 5Figure 5.

Median values of Maximum tensile stresses of specimens with respect to their pause timeThe graph shows a clear difference between maximum tensile stress values of specimens with process interruption (5, 15, 30 and 60 minutes of pause) and specimens without any process interruption (0 minutes of pause). However, there is no clear difference in maximum tensile stress values of specimens with different pause time (>0 minutes).The post-hoc test performed on the data was the Mann Whitney U test, which is a non parametric test to compare two independent groups. Pairwise U and p values for Mann Whitney U tests are shown in Table 3. Any p-value of 0.05 or less has been highlighted to indicate statistical significance in difference of maximum tensile stress values. This test helps in identifying the differences which are significant, which could not be obtained using the earlier Kruskall Wallis H test.854Table 3.

This test helps in identifying the differences which are significant, which could not be obtained using the earlier Kruskall Wallis H test.854Table 3. Post Hoc analysis results of pairwise comparison of different pause time treatment.TreatmentMann Whitney U test0 mins |Mdn=12.9175 mins | Mdn=7.89915 mins | Mdn=7.03730 mins | Mdn=6.13160 mins | Mdn=7.1730 mins |Mdn=12.9175 mins | Mdn=7.89915 mins | Mdn=7.03730 mins | Mdn=6.131U p U p U p U p-0 0.05 - -0 0.05 3 0.35 - -0 0.05 2 0.2 1 0.1 - -0 0.05 4 0.5 2 0.2 0 0.05The Mann-Whitney U test indicated that the ultimate tensile strength was greater for specimens with no pause (Median= 12.917 MPa) than for specimens with 5 minutes pause (Median= 7.899MPa), 15 mins pause (Median= 7.037 MPa), 30 minutes pause (Median= 6.131 MPa), and 60 minutes pause (Median=7.173 MPa). This shows that there is a statistically significant effect of process interruption on the tensile strength of the printed part.

This shows that there is a statistically significant effect of process interruption on the tensile strength of the printed part. Since there is no statistically significant difference in maximum tensile stresses compared pairwise for different pause times (>0 minutes), the statitical analysis suggests that the magnitude of pause time has no significant effect on the strength of the part. However, there is a significant statistical difference between maximum tensile stress values for pause time interval of 30 minutes and that of 60 minutes. This difference is inconsistent with results of other pairs of pause time intervals; more data points are required to confirm the reason for this difference.Another observation from the tensile testing of specimens was made on the location of failure (Figure 6). The layer of failure for each specimen with process interruption was always the layer which was paused (center layer).

The layer of failure for each specimen with process interruption was always the layer which was paused (center layer). The layer of failure for specimens without any process interruption varied across the part. This confirms that process interruption introduces weakness on the layer at which the process was interrupted. This information is important for designing functional components to be manufactured by in situ embedding via AM. The location of embed is advised to be decided by considering the location of layer to be paused.(a)(b)Figure 6.

The location of embed is advised to be decided by considering the location of layer to be paused.(a)(b)Figure 6. Examples of specimen failure after testing with (a) process interruption specimens grouped based on time interval for pause, all failed at the layer which was paused, and (b) specimens with no process interruption, failed at varying layers.8553.3 Summary of Key FindingsTensile testing specimens created to study the effects of process interruption allow the authorsto conclude the following: The Maximum Tensile stress value decreases significantly because of process interruption.The median tensile strength reduced to as low as 48% The Maximum Tensile Stress value has no effect due to time duration of pause ≥ 5 minutes.The test verifies that the presence of a process interruption, irrespective of time duration ≥ 5 minutes, has a negative impact on tensile strength of the part.

For a safety critical part or parts requiring high strength, it is advised that the print shoud be restarted if there is any process interruption.4. EFFECTS OF EMBEDDING WITH AND WITHOUT SHAPE CONVERTERFor an embedded part via AM, the stress concentrations caused due to the presence of cavity for inserts, are expected to reduce the tensile strength of the part. When embedding a component with an irregular top surface, covering the cavity with a shape converter on the paused layer provides support to the resumed layer. Embedding components with flat surfaces may not require shape converters, however, the layer adhesion may reduce because of material transition between the top surface of the embed and the resumed layer.

If the shape converter is of the same material as the print, the authors hypothesize that it should assist in better layer adhesion between paused and resumed layer.To understand the effects of this material transition in different embedding processes on the strength of the part, specimens were prepared with four different treatments: two different embedding methods (with and without shape converters), process interruption, and without any process interruption. The specimens were printed in z-axis orientation on Mark One 3D printer with nylon filament. Specimens with and without any process interruption were designed and prepared similarly to those in Section 3. The two types of specimens for embedding were designed with cavities as shown in Figure 7. The first type of specimen was designed with a cavity of cross section 5mm x 11mm at the center layer, with a depth of 0.5mm, in order to embed a small piece of copper tape.

The first type of specimen was designed with a cavity of cross section 5mm x 11mm at the center layer, with a depth of 0.5mm, in order to embed a small piece of copper tape. The copper tape insert was selected because of its applications in direct writing for in situ embedding. Copper tape also offers a smooth surface for the resumed layer. This allows to compare and evaluate the effects of shape converter, when not used for supporting the resumed layer. The cross section of the cavity was decided after multiple print attempts of different cross sections and their success of completion. The second type of embedded specimen was designed in order to embed with a shape converter. This specimen was also designed with a cavity of cross section 5mm x 11mm at the center of the part. The depth of the cavity was selected as 3mm, in order to embed a printable shape converter.

This specimen was also designed with a cavity of cross section 5mm x 11mm at the center of the part. The depth of the cavity was selected as 3mm, in order to embed a printable shape converter. The shape converter was designed and printed with the same system, parameters, and material as the specimen.856(a)(b)(c)Figure 7. Specimen design for copper tape embed (a), and shape converter embed (b), with the dimensions of shape converter shown in (c).To determine the design clearance for the cross section of the shape converter, a set of shape converter with clearance varying from 0mm to 2 mm in both x and y axis of the build were printed. These shape converters were then inserted in the cavity of designed cross section 5mm x 11mm, as shown in Figure 8. As shown in Figure 8b, the shape converter with a design clearance of 1.5mm in both x and y axes was selected as the most suitable dimension of shape converter for embedding.

As shown in Figure 8b, the shape converter with a design clearance of 1.5mm in both x and y axes was selected as the most suitable dimension of shape converter for embedding. The selection was done on the basis of its ease of fit and low clearance with the walls of the cavity.(a)(b)Figure 8. (a) Initial shape converters with different clearance values varying from 0mm to 2mm in both x and y axis and (b) demonstration that clearance of 1.5mm in both x and y axis was found to fit snuggly in the cavity.To find the clearance for shape converter in the z-direction, another set of shape converters were prepared with same cross sections (9.5mm x 3.5mm) but with different thicknesses (varying from 2mm to 3mm, as shown in Figure 9a) to find its fit in a cavity designed with 3mm depth. The alignment of the top layer of the shape converter was checked after inserting it in the cavity.

The alignment of the top layer of the shape converter was checked after inserting it in the cavity. As shown in Figure 9b, the shape converter with designed z axis clearance of 0.25mm from the cavity aligned with the top layer, and was selected as appropriate z axis clearance for the shape converter. The cavity due to the indented shape converters would not solve the purpose of inserting a shape converter to improve strength and the protruded shape converter would be damaging for the print head if the print is resumed. The aligned specimen was chosen so as to magnify the effects (if any) of the inter-layer bond strength between the top layer of shape converter and the resumed layer. Therefore, the shape converter was designed with a clearance of 1.5 mm on x and y axis each. The857z axis clearance was made 0.25 mm so that the top plane of the shape converter coincides with the paused layer of the part, after being inserted in the cavity.(a)(b)Figure 9.

The857z axis clearance was made 0.25 mm so that the top plane of the shape converter coincides with the paused layer of the part, after being inserted in the cavity.(a)(b)Figure 9. (a) Initial shape converters of same cross sections but different thicknesses values, varying from 0mm to 1mm and (b) demonstration that the shape converter with a design clearance of 0.25mm in z axis was the most aligned with the top layer.After the selection of clearance values, the design of the specimens were printed for testing. The printing parameters and testing environment were the same as in Section 3. The printing process was paused for 15 minutes right after the cavity was printed, to embed the copper tape with dimensions 5mm x 11mm, inside the cavity.4.2 Results and DiscussionsThe statistics of the data obtained for the maximum tensile stresses of the specimens for eachtreatment are reported in Table 4.Table 4.

Tests for NormalityKolmogorov-Smirnova dfStatisticSig.StatisticShapiro-Wilk dfSig.Maximum_Tensile_Stress a. Lilliefors Significance Correction.23312.071.81512.014The Kruskall Wallis H test showed that there was a statistically significant difference in Maximum Tensile stress values between different treatments, χ2(2) = 9.359, p=0.025, with a mean rank of 11.00 for no process interruption, 3.67 for copper tape embed, 3.33 for shape converter embed, and 8.00 for specimens with process interruption. Effect size was calculated as below:858Effect size = χ2/(N-1)=9.359/(12-1)=0.8508The value 0.8508 of effect size means that about 85% of the variability in rank scores is accounted for, by different treatments. Statistical analysis (Mann Whitney U Test) was performed pairwise on the groups of treatments, to find the significance in difference of maximum tensile stress.

Statistical analysis (Mann Whitney U Test) was performed pairwise on the groups of treatments, to find the significance in difference of maximum tensile stress. Pairwise U and p values for the test are shown in Table 6; any p-value of 0.05 or less has been darkened to signify statistical significance in difference of ultimate tensile stress.Table 6.

Pairwise U and p values for the test are shown in Table 6; any p-value of 0.05 or less has been darkened to signify statistical significance in difference of ultimate tensile stress.Table 6. Post Hoc analysis results of pairwise comparison of groups with different embedding conditions, only process interruption, and no process interruption.ConditionNo interruption |Mdn=12.917Process Interruption| Mdn=7.044Copper tape embed| Mdn=5.121Mann Whitney U test U p U p U pProcess Interruption | Mdn=7.044 0 0.05 - -Copper Tape embed | Mdn=5.121 0 0.05 0 0.05 - -Shape converter embed | Mdn=4.976 0 0.05 0 0.05 4 0.827The Mann Whitney U test indicated that the ultimate tensile strength for specimens with embedding (Copper tape Mdn=5.048 MPa & Shape converter embed Mdn=4.976 MPa) is significantly less than for specimens with just a pause (Mdn= 7.173 MPa), U=0, p=0.05. This can be explained by the presence of cavity in the specimens which were embedded.

This can be explained by the presence of cavity in the specimens which were embedded. Theoretically, presence of cavity in any material is responsible for stress concentration around the cavity, making that cross section weaker. In addition, statistical analysis shows that the maximum tensile stress values of embedded specimens (Copper tape Mdn=5.048 MPa & Shape converter embed Mdn=4.976 MPa) are significantly lower than that of specimens with no process interruption (Mdn=12.917MPa), U=0, p=0.05. This difference is higher than the difference observed between process interruption and no process interruption in Section 3.The Mann-Whitney U test also indicated that the ultimate tensile strength for specimens with only copper tape embed (Mdn= 5.121 MPa) and for specimens with shape converter (Mdn= 4.976 MPa) have no statistically significant difference, U= 4, p=0.827.

While the authors hypothesized that shape converters would provide increased adhesion between the last layer before embedding and the resumed layer, there was no improvement in the strength of the parts.The location of failure for the embedded specimen (Figure 10) was same as that for process interruption seen earlier. All the specimens with any form of process interruption failed at the layer which was paused. Though the two types of embedded specimens had two different depths of cavities (0.5mm for copper tape embed and 3mm for shape converter embed), the failure only occurred at the layer paused. This is an important informatrion for making design decisions; the location of the embed in the design of a loaded part becomes crucial because of these two factors.859(a)(b)Figure 10.

This is an important informatrion for making design decisions; the location of the embed in the design of a loaded part becomes crucial because of these two factors.859(a)(b)Figure 10. After tensile testing, (a) copper tape and (b) shape converter embedding specimens with demonstrated failure at the pause layer4.3 Summary of Key FindingsTensile testing specimens created to study the effects of embedding helped the authors toconclude the following: The presence of an embed causes a significant decrease in Maximum Tensile Stress value compared to that with no process interruption. The median stress value reduced to as low as 38% because of the embed. The presence of embed causes a significant decrease in Maximum Tensile Stress value compared to that with only process interruption.

A reduction to 70% of the median stress value was observed. The presence of a shape converter has no significant effect on tensile properties ascompared to an embed without shape converters. The weakness is introduced on the layer which is paused, regardless of the embeddingprocess, and the depth of the embedding cavity.It can be assumed that the process of embedding significantly reduces the strength of the part because of two factors; process interruption and presence of a cavity for embedding. Although one might assume that the use of a shape converter for embedding should assist in improving the adhesion between the layers paused and resumed for embedding, the experiment demonstrated that this is not the case.5. CLOSURE AND FUTURE WORKFor AM embedding applications to effectively produce end-use functional products, design for AM is crucial.

CLOSURE AND FUTURE WORKFor AM embedding applications to effectively produce end-use functional products, design for AM is crucial. Design for AM involves physical design considerations for components to be embedded, deeper understanding of manufacturing process and its effects on the strength of the part. Knowing how and where a process weakens a part is important information for decision- making while designing. This paper investigated two major steps of an embedding process via material extrusion and their impact on part strength; process interruption and embedding with and without the use of a shape converter. It was found that the process interruption introduces a weakness at the layer of pause and reduces the tensile strength of the part, irrespective of the time duration of pause ≥ 5 minutes. The presence of a shape converter while embedding had no impact on the strength of the part compared to embedding without a shape converter.

The presence of a shape converter while embedding had no impact on the strength of the part compared to embedding without a shape converter. However, the embedding process reduced the strength of the part compared to that with just process interruption.860A weakness was introduced on the layer paused, regardless of the depth of cavity used for embedding, which resulted in a consistent failure location. This research work provides first known investigation of mechanical influences of process interruption and embedding process on parts made with material extrusion.These findings provides a wide scope for future work on other in-situ embedding factors, such as the effects of shape converters with different cross sectional areas. A larger set of data for investigating the impact of process interruption with respect to time duration of pause can help in better understanding of abnormality observed in the statistical analysis.

A larger set of data for investigating the impact of process interruption with respect to time duration of pause can help in better understanding of abnormality observed in the statistical analysis. Also, this investigation was done for pause time durations varying from 5 minutes to 60 minutes. An investigation of shorter time durations of process interruption than 5 minutes might show an impact on strength of the part. Future work will aim to expand on the impact of material transition on strength when the print is resumed. A range of material will be investigated for layer adhesion while embedding, so as to provide a database of information about their influences on strength. A range of other process types under AM will also be compared for the influence of embedding on mechanical properties of the parts. This can aid in making better decisions for material and process selection while designing.6.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperInfluence of Gage Length on Miniature Tensile Characterization of Powder Bed Fabricated 304L Stainless SteelS. Karnati*, J. L. Hoerchler*, F. Liou*, J.W. Newkirk†Department of Mechanical and Aerospace Engineering†Department of Materials Science and EngineeringMissouri University of Science and Technology, Rolla, USAAbstractMiniature tensile specimens with varying aspect ratios were fabricated from 304L stainless steel (SS) made using powder bed additive manufacturing (AM) process. The tensile characteristics measured from these specimens were analyzed to assess the impact of gage length. The study found no impact upon varying gage length on yield and ultimate strength measurements. However, a significant impact was observed on strain measurements.

The study found no impact upon varying gage length on yield and ultimate strength measurements. However, a significant impact was observed on strain measurements. This data was also used to perform Weibull statistics to estimate the stochastic performance of the material. Fractography was performed to visually identify the types of flaws. A comparative study with specimens fabricated from cold rolled annealed 304 SS was also performed. The Weibull parameters were used to compare the variability within cold rolled annealed and AM 304L SS. This study indicates miniature tensile testing is a robust characterization technique for obtaining representative material properties.IntroductionOver the past few decades, Additive Manufacturing (AM), more commonly known as 3D printing has gathered a lot of attention from manufacturing and research communities.

The layer by layer build schema of AM provides vast flexibility and the feasibility towards fabricating complex geometries from a multitude of raw material feedstocks. The capabilities offered by AM are expected to significantly alter the existing production infrastructure with respect to production, assembly, and supply chain. Currently, commercial solutions based on metal powder bed processes are available under the names, Selective Laser Melting, Laser Cusing and Direct Metal Laser Sintering (DMLS). These technologies usually employ a high-powered laser to achieve consolidation of metal powder under an inert atmosphere. They can produce parts with a relatively smooth surface finish and lower assembly requirements. In theory, depending on the size of the component and the equipment’s build volume, multiple parts can be built in a single cycle of a powder bed process.

In theory, depending on the size of the component and the equipment’s build volume, multiple parts can be built in a single cycle of a powder bed process. Parts can be laid side by side and stacked on top of each other to attain maximum production output (“Metal Additive Manufacturing Processes” 2017; Gao et al. 2015). Owing to these reasons, coupled with infinite design flexibility and manufacturing ease, the manufacturing industry has shown significant interest towards incorporating powder bed processes into their manufacturing and production lines.©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.289Unlike the parts made from conventional subtractive manufacturing processes likemachining, the properties of AM products can be very different from the feedstock material (Song et al. 2015).

2015). This is due to the fact that the complex mechanics of a powder bed process often leads to very complicated and non-equilibrium outcomes. Evidence suggests that the properties of the fabricated material could even be anisotropic. In general, the performance of powder-bed processed materials can be sensitive to fabrication conditions and are also influenced by important process attributes such as part shape, part size, orientation in the build etc. (Shamsaei et al. 2015; Olakanmi, Cochrane, and Dalgarno 2015). Anisotropies consequent of varying the orientation have been observed by researchers for several materials. Blackwell 2005, found significant differences in tensile properties of IN718L upon varying build orientation. Yadollahi et al. 2017 also found significant differences in tensile and fatigue performances of 17- 4 precipitation hardening stainless steel by varying the build orientation.

Yadollahi et al. 2017 also found significant differences in tensile and fatigue performances of 17- 4 precipitation hardening stainless steel by varying the build orientation. Differences in microstructure and flaw orientations were theorized to cause the varied material performance. Similar effects were noticed by Rao et al. 2016 during the production of the A357 aluminum alloy. The similarities among all the above-discussed works are that the anisotropy and effects of process parameter variation occur at a finer scale during AM production. Assessing the presence and scale of such anisotropy can be infeasible using existing methods that have been standardized for mechanical testing. The sizes of standard test specimens can be too large and thereby be unable to capture and study anisotropy at the desired smaller scales (ASTM E-8, 2016). In this regard, the testing of miniaturized specimens can be an efficient and economical way to characterize such anisotropy in AM materials.

In this regard, the testing of miniaturized specimens can be an efficient and economical way to characterize such anisotropy in AM materials. The smaller size of the specimens makes specimen preparation cheaper and easy. The economy of the approach also makes it feasible to perform extensive rigorous testing and assess stochastic performance easily (Salzbrenner et al. 2017).Miniature testing methodologies, especially mini tensile testing has been a topic of study by researchers of the nuclear industry (N. F. Panayotou, S. D. Artkin, R. J. Puigh 1986; Rosinski et al. 1993). The miniature specimens were used to study irradiated material while ensuring minimal exposure to the operator. Other studies have been performed to identify the impact of specimen design, aspect ratio, thickness to width ratio etc. on property measurements.

Other studies have been performed to identify the impact of specimen design, aspect ratio, thickness to width ratio etc. on property measurements. Researchers have pursued many experimental and theoretical studies to investigate and prove the validity while simultaneously developing mathematical models of the miniature testing methods. The studies concluded that testing miniature specimens can be a valid approach for estimating a material’s performance life and reliably characterize its properties (Kumar et al. 2014; Lord et al. 2010; Dieter 1986). The current study involves the development and implementation of miniature tensile testing methodologies for testing metallic dog-bone shaped specimens. The novelty of the current study includes the specimen size, preparation methodology, and the test specimen’s unique fixturing technique.

The novelty of the current study includes the specimen size, preparation methodology, and the test specimen’s unique fixturing technique. The comparative study between cold rolled and powder bed AM fabricated materials at a multitude of gage lengths is aimed at understanding the property measurements and their relationship to their manufacturing process. The Weibull©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.290statistics provide insights into the variability in AM material in comparison to cold rolled annealed material.Experimental setup and methodTo study the impact of gage length on the tensile property measurements, specimens of different gage lengths were prepared and tested. The dimensions of the mini tensile specimens are shown in Figure 1.

The dimensions of the mini tensile specimens are shown in Figure 1. The gage length of the specimens was varied from 3mm to 6mm in 1mm increments, while the dimensions of the remaining features were kept unchanged. The specimens for testing the powder bed process fabricated material were cut from 4 blocks of 25.4mm (1”) x 25.4mm (1”) x 50.8m (2”) 304L stainless steel. These four blocks of steel were fabricated during a single build with parameter settings optimized for maximum density. The only difference among these blocks was their location on the build plate. A Renishaw AM250 powder bed machine was used to fabricate these blocks while a Sodick wire-EDM machine was used to machine the required specimens. The various specimen designs were cut as columns along the height of the steel blocks. These columns were then sectioned at repeating 1mm thicknesses to obtain the individual specimens. By doing so, the gage length of each specimen was perpendicular to the build direction.

These columns were then sectioned at repeating 1mm thicknesses to obtain the individual specimens. By doing so, the gage length of each specimen was perpendicular to the build direction. The specimens for testing wrought material were cut from a cold rolled and annealed 304 stainless sheet of thickness 1.168mm (0.046”). The gage length of these specimens was aligned along the rolling direction.Gage length, X (mm)3456Figure 1. Drawing of the mini tensile specimen design with the different values of gage length used for this studyTensile testing of these specimens was performed using an Instron universal testing machine. Custom grips were manufactured and setup in order to perform the testing of mini tensile specimens. To avoid any torsion or bending during testing, the grips were designed to be self –aligning. Each self-aligning grip has two joints, which can rotate perpendicular to each other (similar to a ball and socket joint).

Each self-aligning grip has two joints, which can rotate perpendicular to each other (similar to a ball and socket joint). This flexibility was expected to avoid issues of miss- alignment and ensure a uniaxial tensile testing during every test. The setup including the grips©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.291and the extensometer is shown in Figure 2. By using the extensometer to record elongation, closed-loop, strain controlled testing was performed. An extensometer of 15.24 mm (0.6 inches) gage length was used to run these tests. Due to the small size of these specimens, setting the extensometer directly on the specimen was not feasible. The extensometer was setup on the grips of the tensile tester. It was assumed that all the deformation noted by the extensometer came from elongation of the specimen.

The extensometer was setup on the grips of the tensile tester. It was assumed that all the deformation noted by the extensometer came from elongation of the specimen. A strain rate of 0.015mm/mm/min was used leading up to a strain of 0.01. Later, the extensometer was removed and a strain rate of 0.5mm/mm/min was then used to pull the specimen to fracture. The same strain rate schema was used for all the specimens.A detailed statistical analysis was performed to identify the differences among thevarious measurements and also investigate the impact of varying the gage length. The analysis was performed using SAS and JMP software, plots of Weibull distributions were plotted using Minitab software. The estimates for the distributions were calculated using the maximum likelihood methodology. The total number of valid tests at each gage length for both the AM fabricated and cold rolled annealed materials are listed in Table 1.

The total number of valid tests at each gage length for both the AM fabricated and cold rolled annealed materials are listed in Table 1. The fracture surfaces of the broken specimens were imaged using a scanning electron microscope to assess the fracture mechanism and identify the sources of the flaws.Figure 2. Self–aligning grip setup with extensometer attached on the grips (left) and broken mini-tensile specimen in the grips after completion of tensile test (right)©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.292Table 1.

Total number of specimens tested for each type of material at different gage lengths Powder bed 39 34 36 26Gage length 3 4 5 6Cold rolled annealed 30 22 25 27Results and discussionsThe current analysis includes a comparison of 0.2% offset Yield Strength (YS02),Ultimate Tensile Strength (UTS), and Maximum Strain (MxS). The data gathered from specimens made from cold rolled annealed material and powder bed AM material was processed by performing an analysis of variance and Weibull statistical assessment. Box plot charts of these measurements are shown in Figures 3 and 4. At each gage length, the horizontal line in the box represents the median and the box represents the range between the 25th and 75th quantile. The vertical lines represent the range between the minimum and maximum values obtained at the corresponding gage length. Finally, the asterisks represent outlier data points.4203903606543Gage Length (mm)Max.

Finally, the asterisks represent outlier data points.4203903606543Gage Length (mm)Max. strainUTS (MPa)YS02 (MPa)1.11.00.965436543780765750Figure 3. Box plot charts of yield strength, ultimate strength and maximum strain measurements gathered from different gage length specimens of cold rolled annealed stainless steel©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.2936543UTS (MPa)65430.70.60.5540520500Gage length (mm)Max. strain7006756506543YS02 (MPA)Figure 4. Box plot charts of yield strength, ultimate strength and maximum strain measurements gathered from different gage length specimens of powder bed fabricated stainless steelConsidering gage length as the only factor, a one-way factorial analysis was performedon the measurements from both cold rolled annealed and AM material.

The yield strength of the AM material was observed to be higher than the cold rolled annealed material. The higher strength of the AM material can be attributed to the smaller grain size that is characteristic of the AM process. However, the ultimate tensile strength and the maximum strain values of the cold rolled annealed material were higher than the AM material. The analysis of variance revealed there were no significant differences in the yield strength and ultimate strength measurements for tests performed on the AM material. However, with varying gage lengths, a significant difference was observed among the maximum strain measurements. The same analysis on cold rolled annealed material measurements concluded there were no significant differences among the ultimate strength measurements. However, in this case, the yield strength and maximum strain measurements showed a significant difference.

However, in this case, the yield strength and maximum strain measurements showed a significant difference. The statistical significance of these differences was also validated by running a Tukey’s analysis of means comparison.Unlike the AM material, the yield strength values of the cold rolled annealed materialwere observed to be varying with gage length. The authors theorize that this variation could be related to the larger grain size in the cold rolled annealed material. Due to the larger grain size, the authors theorize that the gage lengths investigated might not be voluminous enough to be representative of the cold rolled annealed material.

Due to the larger grain size, the authors theorize that the gage lengths investigated might not be voluminous enough to be representative of the cold rolled annealed material. With increasing gage length, the probability©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.294of the number of larger grains within the gage length is expected to increase and thereby cause a drop in yield strength. However, the drop was not a substantial amount (10 MPa approx.) while still being statistically significant. The values of yield strength (350 MPa) and ultimate strength (690 Mpa), as noted in the certificate of test supplied with the material, were lower than the values obtained through mini tensile testing. This supports the claim that these cold rolled annealed specimens were not representative of the entire material.

This supports the claim that these cold rolled annealed specimens were not representative of the entire material. The higher ultimate strength could be attributed to the differences in flaw population between mini tensile specimens and standard tensile specimens. The smaller gage volume of the cold rolled annealed specimens would also have a fewer number of flaws. Therefore, these specimens could perform better during mini-tensile testing which is reflected in the larger measurements of the ultimate strength values.The variability in property measurements of AM material was much higher than the variability in measurements from the cold rolled annealed material. The yield and ultimate tensile strength measurements on the cold rolled annealed material were very consistent and were spread over a small range (approximately 20 MPa) in comparison to AM property measurements (60-80 MPa).

The wide variability in AM property measurements was suspected to be from the differences in the steel blocks used to prepare the specimens.BlockMax. StrainUTS (MPa)YS02 (MPa)72069066043210.80.60.443215604804004321Figure 5. Yield strength (MPa), ultimate strength (MPa) and maximum strain measurements from specimens with a gage length of 3mm prepared from different blocks.©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.295As power and other fabrication, parameters were kept the same, and all the blocks werebuilt during a single run of the machine, the only possible source of variation was expected to be the location of these blocks on the substrate build plate.

The analysis of variance was then re- investigated with gage length and the source material (as a blocking factor to isolate the variation at the source) as the factors of variation. Based on the new analysis of variance, the different source materials were confirmed to be the cause of this high variability. The yield strength, ultimate strength and maximum strain measurements of specimens with 3mm gage length gathered from the four different blocks are shown in Figure 5. The yield and ultimate strength values of specimens from block 2 were found to be higher than the specimens from the remaining blocks. The maximum strain values from specimens of block 4 were much lower than the specimens from the remaining block. For the sake of brevity and clarity, the sources of these variations are discussed in detail in the latter parts of this discussion.

For the sake of brevity and clarity, the sources of these variations are discussed in detail in the latter parts of this discussion. Also, by grouping the measurements by the source material, it can be noticed that the spread in mechanical properties of material form individual blocks of AM material is similar to that of cold rolled annealed material.The values of maximum strain were noticed to vary significantly corresponding to thevarying gage lengths for both cold rolled annealed and AM materials. It is well established that varying gage length produces different measurements of ductility. Also, geometrically similar specimens are expected to yield similar measurements. However, the similarity only minimizes the difference but does not eliminate it. The maximum strain measured during tensile tests is comprised of the elastic deformation, elongation from uniform plastic deformation (volume constancy relation holds true) and localized deformation i.e. necking.

necking. The amount of elastic deformation is negligible in comparison to the plastic deformation. The strain from uniform plastic deformation would be similar for specimens of all gage lengths as it is dependent on the material properties and not the gage length. The elongation from necking is also expected to be the same for all gage lengths as, all the specimens have the same cross section (nominal 1mm x 1mm). However, as the gage length increases the contribution of elongation from necking to the maximum strain value decreases (necking elongation is divided by larger gage length values). This decrease is seen in the maximum strain values of both AM and cold rolled annealed materials.

This decrease is seen in the maximum strain values of both AM and cold rolled annealed materials. The total elongation after neglecting elastic deformation can be represented in relation to the gage length as written below𝑢𝑇 = 𝐺𝐿 ∗ 𝑒𝑢 + 𝛼Where 𝑢𝑇 is the total elongation, GL is the gage length, 𝛼 is the elongation from necking and 𝑒𝑢is the strain corresponding to uniform plastic deformation.In order to decouple strain measurements and estimate the strain corresponding to uniform plastic deformation and localized deformation, a regression analysis was performed. The analysis was performed to establish the linear empirical relationship between the total elongation©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.296and the nominal gage length of the specimens.

The elongation values calculated from the maximum strain measurements were fit against the nominal gage lengths. The SAS software was used to perform these linear fits. The intercept of the linear fit would be an estimate of the contribution in elongation from localized deformation. The slope of the linear fit would be an estimate for the strain that corresponds to uniform plastic deformation. The estimated values of slope and intercept from the linear fits for both cold rolled annealed and AM materials are listed in Table 2. The higher values of slope and intercept of the cold rolled annealed material imply that cold rolled annealed material is better than AM material in terms of ductility. The slope value for cold rolled annealed material implies that the cold rolled annealed material can undergo larger uniform plastic deformation than AM material. The higher value of intercept suggests that cold rolled annealed material necks larger than the AM material.

The higher value of intercept suggests that cold rolled annealed material necks larger than the AM material. The higher standard error values for slope and intercept estimates of AM material implies higher variability in ductility and lower consistency in performance when compared to the cold rolled annealed material (as seen from Figures 6 and 7). This also concludes the cold rolled annealed material is better than AM material in terms of toughness (area under the stress strain curves as seen from Figures 6 and 7).Table 2. Slope and intercept values of linear fit between total elongation and nominal gage lengthCold rolled annealed materialAM materialVariableParameter EstimateStandard ErrorParameter EstimateStandard Error𝛼1.216mm0.0230.545mm0.129𝑒𝑢0.6880.0050.4670.029Figure 6.

Engineering stress strain curves for mini tensile specimens of 3mm gage length made from AM 304L stainless steel©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.297Figure 7. Engineering stress strain curves for mini tensile specimens of 3mm gage length made from cold rolled annealed 304L stainless steelEmpirical cumulative probability distributions of yield strength, ultimate strength, and maximum strain were plotted using two-parameter and three-parameter Weibull distributions. The American society of testing and materials standards C1683 and C1239 outline the basis and methodology for performing statistical analysis on material properties of advanced ceramic materials. However, the Weibull distribution has also been used to model the stochastic performance of metals (S. Karnati et al. 2016; J.W. Newkirk, T. Amine, and J. Koth 2016; J.W.

However, the Weibull distribution has also been used to model the stochastic performance of metals (S. Karnati et al. 2016; J.W. Newkirk, T. Amine, and J. Koth 2016; J.W. Newkirk and J. Wang 2014; Salzbrenner et al. 2017). The equations of two-parameter and three- parameter Weibull distributions are as follows,𝑃(𝑥) = (𝑚 𝛽) (𝑥 𝛽𝑚−1 )𝑒𝑥𝑝 [− (𝑥 𝛽)𝑚] , 𝑥 > 0𝑃(𝑥) = (𝑚 𝛽) (𝑥 − 𝜃 𝛽𝑚−1 )𝑒𝑥𝑝 [− (𝑥 − 𝜃 𝛽𝑚 )] , 𝑥 > 0, 𝜃 > 0Where P is the probability of event the occurring, x is the random variable, m is the Weibull modulus or shape parameter, β is the scale or characteristic value and θ is the threshold value. Weibull modulus or the shape parameter describes the breadth of the distribution. Higher the value of the Weibull modulus, lower is the spread of the distribution. The three parameter Weibull distribution can be used to model data where the possibility of failure under a certain value of the random variable has zero probability.

The three parameter Weibull distribution can be used to model data where the possibility of failure under a certain value of the random variable has zero probability. The threshold value is an estimate of the value under which the probability of failure occurring is equal to zero.The yield strength values from both materials were found to be a better fit for three- parameter Weibull distributions. The ultimate strength and maximum strain values were fit using a two-parameter Weibull distribution. The type of fit was chosen basing on the Akaike and©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.298Bayesian information criterion provided by JMP software. The Weibull fits were performed using JMP software by following the maximum likelihood approach. The plots of these fits were made using Minitab software.

The Weibull fits were performed using JMP software by following the maximum likelihood approach. The plots of these fits were made using Minitab software. The Weibull distribution plots for yield strength and 95% confidence intervals for the threshold, Weibull modulus, and characteristic strength were as shown in Figures 8, 9 and 10. The 95% confidence intervals of the distribution parameters for ultimate strength and maximum strain measurements and corresponding Weibull distribution plots are shown in Figures 11, 12, 13, 14 and 15.Figure 8. Cumulative three-parameter Weibull fits of the yield strength data obtained from specimens of SLM and cold rolled annealed (CRA) materials. The gage lengths are 3 (blue), 4 (red), 5 (green) and 6 (purple)8055060500402045004003456Gage lengthCharacterisitic strengthWeibull modulusThresholdFigure 9.

95% confidence intervals of two-parameter Weibull fit (see Figure 15) on ultimate strength data of cold rolled annealed material. Weibull α is the characteristic strength and Weibull β is the Weibull modulusn a r t si1.51150100c i t s i r e t c a r a h C0.503456500Gage lengthWeibull αWeibull βFigure 14. 95% confidence intervals of two-parameter Weibull fit (see Figure 15) on maximum strain data of cold rolled annealed material. Weibull α is the characteristic strain and Weibull β is the Weibull modulusThe Figures 10-17 give a good picture of the tensile stochastic performance of the cold rolled annealed and powder bed AM materials. The previously identified decreasing trends in yield strength measurements for cold rolled annealed material and maximum strain measurements of both cold rolled annealed and AM materials can also be seen from these Weibull statistics.

The decreasing trend in threshold values of yield strength for cold rolled annealed materials indicates the drop in measured values of yield strength with increasing gage length. The Weibull moduli of both AM and SLM materials are comparable and do not vary significantly with increasing gage lengths.The characteristic strain values for both cold rolled annealed and AM material indicate a drop in the measured values with increasing gage length. The Weibull modulus of strain for the©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.301s u u d o Mll lu b e Wis u u d o Mll lu b e Wicold rolled annealed material is substantially higher than that of AM material. This is indicative of the better performance of the cold rolled annealed material and the large variability in ductility of the AM material.

This is indicative of the better performance of the cold rolled annealed material and the large variability in ductility of the AM material. The same conclusions apply to the ultimate strength data as well. For some of the cases, the 95% confidence intervals were very wide and have significant overlap. A larger sample size can help decrease the error and provide better estimates to help resolve these overlaps. At this point, no change is observed in Weibull statistics with increasing gage length.Figure 15. Two-parameter Weibull distributions of ultimate strength (UTS) and maximum strain (MxS) for cold rolled annealed and SLM/AM material. The different gage lengths are 3 (blue), 4 (red), 5 (green) and 6 (purple)The lower performance of the AM material in terms of ultimate strength and ductility wasinvestigated by analyzing the images of all the fracture surfaces gathered from the broken specimens of 3mm gage length.

The specimens yielding high values of yield strength and ultimate strength, in general, had minimal porosity. The total porosity was made up of gas (spherical) and lack of fusion (irregular and along the direction of build) porosities. Specimens with lower strengths were observed to contain pores on the fracture surface.The fracture surfaces of specimens yielding highest and lowest values of strength andductility are shown in Figure 16. The specimen that yielded the highest strength values was©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.302observed to have minimal porosity (Figure 16a). However, the specimen that yielded the highest strain (Figure 16b) was observed to have spherical pores on the fracture surface. The authors theorize this discrepancy can be attributed to the strain component stemming from localized deformation i.e. necking.

The authors theorize this discrepancy can be attributed to the strain component stemming from localized deformation i.e. necking. The specimen in Figure 16a has a larger cross sectional area in comparison to the specimen in Figure 16b. This is indicative of necking to a larger extent in specimen shown in Figure 16b. The difference in elongation from necking produces a higher strain measurement and is therefore not a true representation of the specimen ductility.(a)(b)(c)Figure 16.

The difference in elongation from necking produces a higher strain measurement and is therefore not a true representation of the specimen ductility.(a)(b)(c)Figure 16. Secondary electron images of the fracture surfaces of (a) a specimen from block 2 with highest measured values of yield and ultimate strength (b) a specimen from block 1 with highest measured value of maximum strain and (c) a specimen from block 4 with the lowest measured values of yield strength, ultimate strength and maximum strain©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.303By this reasoning, the lower performance of strain in specimens from block 4 could possibly be attributed to the difference in necking elongation. However, this needs a more detailed study into the reasoning behind the early fracture of these specimens.

However, this needs a more detailed study into the reasoning behind the early fracture of these specimens. The specimens that yielded the lowest values of strength and strain was observed to have a lack of fusion pores across the cross section of the specimen.ConclusionsThe current effort involved the fabrication and testing of mini tensile specimens ofdifferent gage lengths from conventional and powder bed AM fabricated 304 SS. To assess the impact of gage length, the measured values of 0.2% offset yield strength, ultimate tensile strength and maximum strain were analyzed and the following conclusions were madeAnalysis of variance revealed a significant impact of gage length on the maximum strain measurements. With the increase in gage length, the maximum strain values were observed to decrease.

With the increase in gage length, the maximum strain values were observed to decrease. This was attributed to the decreasing contribution of localized deformation to maximum strain valuesA regression analysis was performed to estimate the strain corresponding to uniform plastic deformation and localized deformation. The analysis revealed the cold rolled annealed material was more ductile and tougher than the powder bed AM material  The yield strength of powder bed AM material was observed to be higher than the cold rolled annealed material. However, the ultimate strength and maximum strain values of AM material were lower than those of cold rolled annealed materialThe variability in powder bed AM material was observed to be higher than the cold rolled annealed material.

The wide variability in AM material was identified to be associated with the differences in the source material used to fabricate the different mini tensile specimensWeibull statistics were also performed on these tensile property measurements. The estimates from the Weibull fits were in accordance with the conclusions from analysis of variance.The Weibull estimates for the yield strength of cold rolled annealed and AM material were comparable and has similar Weibull moduli. The estimated values of Weibull moduli for ultimate strength and maximum strain also corroborated the wide variance in performance of powder bed AM material Investigation of fracture surfaces revealed porosity was the cause for such wide variabilityFinally, the mini tensile specimens were able to produce repeatable and reliable measurements of the tensile properties of conventional and powder bed fabricated AM materials.

These specimens were also able to capture the source to source variation in terms of strength and ductility©2017 The Department of Energy’s Kansas City National Security Campus is operated and managed by Honeywell Federal Manufacturing Technologies, LLC under contract number DE-NA0002839.304AcknowledgmentsThis work has been funded by Honeywell Federal Manufacturing & Technologies under Contract No. DE-NA0002839 with the U.S. Department of Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United States Government retains a nonexclusive, paid up, irrevocable, world-wide license to publish or reproduce the published form of this manuscript, or allow others to do so, for the United States Government purposes.Furthermore, the authors would like to acknowledge Kyle Stagner, Ian Wille, and BrianBullock for their aid in this research.ReferencesBlackwell, P.L. 2005.

lid Freeform Fabrication 2018: Proceedings of the 29th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperINFLUENCE OF GRAIN SIZE AND SHAPE ON MECHANICAL PROPERTIES OF METAL AM MATERIALSR. Saunders†, A. Achuthan‡, A. Iliopoulos†, J. Michopoulos†, and A. Bagchi††Materials Science and Technology Division, U.S. Naval Research Laboratory, Washington, DC, 20375 ‡Dept. of Mechanical and Aeronautical Engineering, Clarkson University, Potsdam NY, 13676AbstractMetal powder-based additive manufacturing (PAM) typically results in microstructures with a texture and columnar grain structure. The columnar grains can vary greatly in size and shape throughout the microstructure, which can significantly affect the mechanical properties of the resulting part.

The columnar grains can vary greatly in size and shape throughout the microstructure, which can significantly affect the mechanical properties of the resulting part. A previous study developed a microstructurally informed crystal plasticity con- stitutive model that took into account grain sizes and shapes then showed that grain geometry can influence the prediction of mechanical behavior of the part. In the present work, the influence of grain aspect ratio, size, and loading direction on the resulting mechanical properties of the PAM part are investigated through a parametric study. Results show that considering size and shape effects have the tendency to increase the material yield strength while decreasing the initial strain hardening modulus.

Results show that considering size and shape effects have the tendency to increase the material yield strength while decreasing the initial strain hardening modulus. Using this knowledge, it may be possible to optimize a PAM microstructure using process parameters to produce a part which exhibit superior yield strength and hardening modulii compared to traditional materials.IntroductionMetal powder-based additive manufacturing (PAM) processes involve continued melt- ing and solidiﬁcation of metallic p owder. The rate at which material cools after solidiﬁcation as determined by the thermal gradient, inﬂuences t he evolution o f t he l ocal m icrostructure. Large thermal gradients lead to substantial local stresses in the solidiﬁed m aterial [ 1]. B ecause o f the elevated temperature, the yield stress of the material is substantially decreased meaning that the local thermal stress can lead to local yielding of the material.

B ecause o f the elevated temperature, the yield stress of the material is substantially decreased meaning that the local thermal stress can lead to local yielding of the material. Local yielding results in non-uniform plastic strain which, when coupled to the mechanical constraint of the support structure (e.g. the base plate), may cause large residual stress and undesirable distortion of the build part. Residual stresses in the build part have been suggested to deteriorate fatigue strength [2] and decrease PAM part’s service life. Additionally, the stress evolution in the microstructure may produce defects, such as dislocations, pores, and microcracks [3].The microstructure resulting from a PAM process typically has a texture due to epitaxial grain growth. The grains are often columnar with different grain sizes and aspect ratios, which affect the mechanical properties of the material. To numerically study these effects, a number ofThis material is declared a work of the U.S.

To numerically study these effects, a number ofThis material is declared a work of the U.S. Government and is not subject to copyright protection in the United States. Approved for public release; distribution is unlimited.1751ﬁnite element (FE) models have been developed. Early models primarily focused on challenges associated with a moving heat source, material evolution, and computational cost of the simulation [4, 5, 6, 7, 8, 9]. Subsequent studies focused on residual stress evolution and the role of process parameters in the PAM process in inﬂuencing these residual stress [10, 11, 12, 13, 14, 15]. A more recent study [16] focused on the development of a microstructure-informed constitutive model de- veloped to describe the mechanical behavior of solidiﬁed material produced by PAM by extending the recently reported [17, 18] microstructural-feature size-dependent crystal plasticity constitutive model for nickel superalloy materials.

A follow-up study [19] developed a method to better cap- ture the grain boundaries, and therefore grain shape, to more accurately represent the local stress intensiﬁcation of irregularly shaped grains.In this work, the models developed previously [16, 19, 20] are extended to examine the effects of grain size and shape on the prediction of the mechanical behavior of a material. First, an overview of the crystal plasticity model along with its implementation is presented. Next, synthetic microstructures that contain features seen in PAM microstructures are generated and meshed to create an FE model of a representative volume element (RVE). The generated FE models are used to conduct a parametric study where average RVE aspect ratio and grain size are varied. As part of this study, the constitutive response is varied to consider cases with no size or shape effects, size effects only, and both size and shape effects.

As part of this study, the constitutive response is varied to consider cases with no size or shape effects, size effects only, and both size and shape effects. The effect of load direction is also examined by considering longitudinal (i.e., along the dominant grain direction) and transverse (i.e., out of plane of the dominant grain direction) tension and shear.Constitutive Model and HomogenizationThe crystal plasticity constitutive model implemented in this work is an extension of the grain-size dependent crystal plasticity constitutive model described in [17, 18]. The constitutive model has been extended by [16] to account for the aspect ratio of grains and is based on a core and mantle framework to describe the effect of a grain boundary. In this framework, a grain boundary inﬂuence region where there is an increased resistance to dislocation nucleation is considered.

In this framework, a grain boundary inﬂuence region where there is an increased resistance to dislocation nucleation is considered. The resistance to dislocation nucleation in the region has a maximum value at the grain boundary and fades away at the inner boundary of the region in the grain. The resistance in the grain bound- ary inﬂuence region is represented in a fasion similat to that of work-hardening, where the grain boundary effects are associated with an increase in strength and decrease in initial strain-hardening modulus. By treating the grain boundary inﬂuence similar to work-hardening, the crystal plasticity constitutive modeling framework used to capture work-hardening effects can be implemented.

By treating the grain boundary inﬂuence similar to work-hardening, the crystal plasticity constitutive modeling framework used to capture work-hardening effects can be implemented. A brief overview of the framework is presented here to elucidate how the grain boundary effects are incorporated; further details can be found in [16].The instantaneous shear strength of a slip system, α, at a material point at location r inthe grain boundary inﬂuence region can be stated as the additive decomposition,g(α)(r, t) = g(α)0 + g(α)GB(r) + g(α)L (r, t),where g(α)0is the shear strength of the material points outside the grain boundary inﬂuence regionThis material is declared a work of the U.S. Government and is not subject to copyright protection 1752 in the United States.

Government and is not subject to copyright protection 1752 in the United States. Approved for public release; distribution is unlimited.(1)of an annealed sample, g(α) L (r, t) rep- resents the increase in strength due to strain-hardening as a result of mechanical loading, and r and t are location and time, respectively. The resistance to dislocation motion due to the grain boundary effect has no evolution in time. Thus, it can be represented as an integration in the total strain in all slip systems, dγ, according toGB(r) represents the grain boundary effect on the strength, g(α)g(α) GB(r) =(cid:90) ˜γGB(r)˙g(α)(˜γ, ˙γ(β))dγ,0where ˙g(α) is the rate of strain hardening, β is any slip system, ˙γ is the shear ﬂow rate, and γ is the cumulative shear ﬂow strain. The grain boundary effects and the effects from loading, as stated above, appear the same as a work hardening type behavior.

The grain boundary effects and the effects from loading, as stated above, appear the same as a work hardening type behavior. However, the grain boundary effect is not a physical quantity but rather a mathematical one introduced to allow for a variation in strength and initial strain hardening modulus. As such, this quantity must be separated appropriately from the actual strain caused by loading and is denoted with the ˜() symbol.

As such, this quantity must be separated appropriately from the actual strain caused by loading and is denoted with the ˜() symbol. Equation 2 can be simpliﬁed by use of the strain-hardening modulus, hαβ, so thatg(α) GB(r) =(cid:90) ˜γGB(r)Σ(β)hαβ(˜γ) (cid:12)(cid:12) ˙γ(β)(cid:12)(cid:12) dγ.0Equation 3 can be further simpliﬁed by considering the special case of a hyperbolic secant squared type strain-hardening with(cid:12) (cid:12) (cid:12) (cid:12)(cid:12) (cid:12) (cid:12) (cid:12)h0∞γ τs − τ0∞hαβ = qαβh0∞sech2,where qαβ differentiates latent-hardening (α (cid:54)= β) and self hardening (α = β), h0∞ is the ini- tial hardening modulus, and τs is the maximum resistance to shear ﬂow. The previous equations are very general since they are deﬁned on a material point basis. However, the computational implementation of the equations on a material point basis can be challenging for a realistic 3D microstructure.

However, the computational implementation of the equations on a material point basis can be challenging for a realistic 3D microstructure. In this work, an alternative approach using homogenization on a grain-by-grain basis is utilized. The total strength can be rewritten as,g(α)(r, t) = τo∞ + gGB +(cid:90) t˙g(α)dt,0wheregGB = g(α)GB(˜γGB)and˜γGB =(cid:82)VGB˜γGB(r)dVV.The above integration is aided by mapping the arbitrary grain geometry to a simpliﬁed domain, such as a sphere or an ellipsoid. Considering only size effects, a simple sphere with the volume equivalent of the grain is sufﬁcient to represent the grain. The diameter of this sphere is found as d = V , where d is the diameter of the representative sphere and V is the known grain volume 3 π) ( 4of the arbitrarily shaped grain. Considering both the size and shape effects requires the use of anThis material is declared a work of the U.S.

Considering both the size and shape effects requires the use of anThis material is declared a work of the U.S. Government and is not subject to copyright protection 1753 in the United States. Approved for public release; distribution is unlimited.(2)(3)(4)(5)(6)(7)ellipsoid, which can be represented by a general quadric surface ψ(x, y, z) = 0, where ψ(x, y, z) is a quadratic polynomial given by,ψ(x, y, z) =A1x2 + A2y2 + A3z2+ A4xy + A5yz + A6xz+ A7x + A8y + A9z + A10.Individual grains are then represented using an ellipsoid that provides the best ﬁt in terms of min- imum linear least square (llsq) error. The approach essentially uses all the known points on the arbitrary grain surface to determine the appropriate coefﬁcients, Ai, for the given grain.

The approach essentially uses all the known points on the arbitrary grain surface to determine the appropriate coefﬁcients, Ai, for the given grain. With a known integrable grain geometry, the integration in Equation 7 can be performed by assuming a constant grain boundary thickness, δGB, and parameterizing the ellipsoid and grain boundary in terms of the radii.Model GenerationIn this work, RVEs of microstructures that are similar in nature to those identiﬁed experi- mentally via Electron Back Scattering Diffraction (EBSD) imaging in actual PAM microstructures are synthesized using a continuum diffuse interface model [20]. The resulting synthetic microstruc- ture generated by the continuum diffuse interface model is represented in ﬁgure 1. The data in the RVE is in the form of a 3D matrix with each (i, j, k) location containing a grain label from 1 to N , where N is the total number of distinct grains in the microstructure.

The data in the RVE is in the form of a 3D matrix with each (i, j, k) location containing a grain label from 1 to N , where N is the total number of distinct grains in the microstructure. The structure of this data is similar to the data structure used in gray scale images such as those obtained from magnetic resonance imaging (MRI) and computed tomography (CT) scans. As such, the voxel image data is imported into ScanIP, a software developed by Simpleware (Synopsys, Mountain View, USA) to semi-automatically segment such image data.Figure 1: RVE of synthetic microstructure generated by the 3-D continuum diffuse interface model.The segmentation process creates a surface representation of each grain, which is then used to generate a surface and volume FE mesh. To eliminate the “stair stepping” effect inherent to the data structure, an algorithm to anti-alias and smooth the data is applied to ensure smoothThis material is declared a work of the U.S.

To eliminate the “stair stepping” effect inherent to the data structure, an algorithm to anti-alias and smooth the data is applied to ensure smoothThis material is declared a work of the U.S. Government and is not subject to copyright protection 1754 in the United States. Approved for public release; distribution is unlimited.(8)contours along the grain boundaries. Following the smoothing of the surface data, the segmented boundaries are converted into a triangulated surface representation, then a multi-part surface dec- imation algorithm followed by a Delaunay tessellation with tetrahedral elements is applied [21].

A tetrahedral mesh is used in this work as it better represents the grain geometries (and therefor overall RVE behavior) at a similar or lower computational cost compared to a hexahedral mesh [19].The ﬁnite element model in this work is implemented in Abaqus/Standard (Dassault Systems, Providence, RI, USA) implicit FEA package [22], with the crystal plasticity constitutive model highlighted earlier, implemented through a user material subroutine (UMAT). The RVE faces have periodic boundary conditions applied. The constitutive model parameters used in this work are taken from a copper cubic lattice structure and from literature [23, 24]. The copper cubic lattice structure used to deﬁne the elastic material properties C11, C12, and C44 and the plastic material parameters ˙a, n, and q were obtained from [23]. The parameters ˙a and n are the parameters for a power law ﬂow rule. The material properties τ0∞, τs and h0∞ (eq.

The parameters ˙a and n are the parameters for a power law ﬂow rule. The material properties τ0∞, τs and h0∞ (eq. 4) were determined by matching the stress-strain behavior of a grain-size independent simulation with the stress-strain behavior reported for the sample with large average grains [24]. A grain boundary thickness δGB of 0.333 µm was chosen. All the material parameters used in this study are shown in Table 1.Table 1: Material parameters used for the numerical simulation.Parameters Value τ0∞ 9 MPa τs 95 MPa ˙a 0.001 /s n 10 q 1 C11 168.4 GPa C12 121.4 GPa C44 75.4 GPa ˜γ∗ 1.07 GB h0∞ 240 MPaNote that the value of n obtained from the previously referenced works [23, 24] was 100, but a value of 10 is used in this study. The parameter n controls the rate dependence of the material and higher values of n result in a more intensive computational problem.

The parameter n controls the rate dependence of the material and higher values of n result in a more intensive computational problem. To demonstrate the effect of the n variability, identical problems were run with only the value of n varying from 10 to 100 (Figure 2). Figure 2a shows how the choice of n inﬂuences the predicted stress-strain behavior and Figure 2b shows how n inﬂuences computational time. The stress-strain behavior is relatively unaffected except for the absolute value of the predicted stress. However, the computational time using n = 10 compared to using n = 100 is increased by a factor of 4. The goal of this study is to generate large amounts of comparative data, thus n = 10 was chosen to expedite the data col-This material is declared a work of the U.S. Government and is not subject to copyright protection 1755 in the United States. Approved for public release; distribution is unlimited.lection process.

Government and is not subject to copyright protection 1755 in the United States. Approved for public release; distribution is unlimited.lection process. The results of the study are not affected by this choice as the material parameters are irrelevant as long as they are consistent between analyses. The results only show changes in predicted mechanical behavior due to changes in how the grains in the RVE are represented, i.e. no size or shape effect, size effect (grains represented by a sphere), or size and shape effect (grains represented by an ellipsoid).

no size or shape effect, size effect (grains represented by a sphere), or size and shape effect (grains represented by an ellipsoid). The trends seen between runs due to changing grain size/aspect ratio and load direction are being analyzed in this purely computational study.(a) Variation of stress and strain with a changing n value.(b) Computational cost in execution time, in hours, to run the example problem as a function of the parameter n.Figure 2: The effect of varying n on mechanical behavior prediction and computational time.Results and DiscussionThree RVE microstructures are generated and analyzed in the parametric study presently being conducted. The three RVEs are generated by using three different combinations of parame- ters in the continuum diffuse interface model. The RVEs are referred to by the anisotropy factor κ0 that controls the anisotropy associated with the grain aspect ratios as described in [20].

The RVEs are referred to by the anisotropy factor κ0 that controls the anisotropy associated with the grain aspect ratios as described in [20]. Three κ0 values of 1, 3, and 5.5 were chosen to generate RVEs with three distinctly different aspect ratios as shown by Figure 3. The three κ0 values were chosen to simulate an equiaxial grain structure (κ0 = 1) such as that seen in the traditional manufacturing process, a slightly directionally-biased grain structure (κ0 = 3), and a highly elongated grain structure (κ0 = 5.5). The last two are representative of features that can be seen in PAM microstructures. Each RVE has approximately 300 grains and a volume of 0.1 mm3.The variation of grain size and shape throughout each of the RVEs is shown in Figure 4, where the grain effective diameter is calculated by equating the grain volume to a sphere and the effective aspect ratio is the largest axis of the ﬁtted ellipsoid to the mean of the intermediate and smallest axes.

From Figure 4a, it can be seen that the number of grains and effective diameter of the grains in each microstructure are approximately the same. This indicates that the three RVEs contain similar number of grains with similar volumes, thus the largest variability in the RVEs are the shapes of the grains. The effective aspect ratios of Figure 4b show the expected behavior that asThis material is declared a work of the U.S. Government and is not subject to copyright protection 1756 in the United States. Approved for public release; distribution is unlimited.(a) κ0 = 1.(b) κ0 = 3.(c) κ0 = 5.5.Figure 3: Synthetic microstructure RVEs generated by the continuum diffuse interface model with three different values of κ0.κ0 is increased the variability in the aspect ratio is increased. However, the increase in κ0 does not mean that grains with a low aspect ratio are completely removed, the inverse is also true.

However, the increase in κ0 does not mean that grains with a low aspect ratio are completely removed, the inverse is also true. This is easily demonstrated by noting that the highest effective aspect ratio of the three RVEs is actually in the κ0 = 3 case and not the κ0 = 5.5 case, but on average the effective aspect ratio in the κ0 = 5.5 case is higher than the κ0 = 3 case.(a) Grain effective diameter (size) distribu- tion sorted by size.(b) Effective grain aspect ratio (shape) compared to the effective grain diameters (size).Figure 4: Synthetic microstructure RVE grain size and shape distributions.The generated RVEs are now used in the previously described parametric study. The goal of this study is to determine how the grain boundary effect changes the stress-strain behavior of the three synthetic RVEs under four different loading conditions.

The goal of this study is to determine how the grain boundary effect changes the stress-strain behavior of the three synthetic RVEs under four different loading conditions. The RVEs are subjected to a uniaxial tensile loading in the dominant grain direction (Y direction in ﬁgure 3) and a direction transverse to this direction (Z direction in ﬁgure 3). An in-plane (Y − Z direction in ﬁgure 3) and an out-of-plane (X − Z direction in ﬁgure 3) shear loading were also considered. Here we have assumed that loading in the X and Z directions and shear in the X − Y and Y − Z planes produces identical behavior. The loading is applied via a speciﬁed displacement to generate aThis material is declared a work of the U.S. Government and is not subject to copyright protection 1757 in the United States. Approved for public release; distribution is unlimited.maximum volume-averaged nominal strain in the direction of the load of 60%.

Approved for public release; distribution is unlimited.maximum volume-averaged nominal strain in the direction of the load of 60%. The analysis results were post-processed by volume averaging the stress and strain in the direction of the load. Other components of stress and strain not in the direction of the applied displacement were nonzero to maintain equilibrium but are not considered here. The cumulative results of the parametric study are shown in Figure 5. Note that these results include data up to 30% strain, as after that point all results have reached the saturation stress, thus coincided with one another and do not contribute to the discussion.The ﬁrst observation that can be noted is that by not including size and/or aspect ratio (AR) effects, the yield point for all values of κ0 is unchanged for the different loading directions, as expected.

Furthermore, the yield point for similar loading (e.g., tension or shear) when not considering size and/or shape effects is also unchanged. Intuitively, this is an unexpected behavior because it is well known that materials with a dominant material orientation such as those seen in PAM-produced parts, generally have an anisotropic response (e.g., carbon ﬁber-epoxy composites and biological materials such as muscle). This reinforces the need for a constitutive model that incorporates grain size and shape effects. Next, these results show that in tension an increasing RVE aspect ratio (controlled by κ0) has the effect of decreasing the stress at 30% strain but in shear, has the effect of increasing the stress at this point. This effect is independent of the grain boundary effect. Including the size effect increases the yield strength and decreases the strain hardening modulus.

This effect is independent of the grain boundary effect. Including the size effect increases the yield strength and decreases the strain hardening modulus. This effect is ampliﬁed when considering the size and shape effects and further ampliﬁed when the microstructure average AR is increased. Due to the RVE grain sizes being the same, on average, the size effect does not vary with varying κ0. Finally, it can be seen that in all cases, the yield point is heavily inﬂuenced by grain shape. However, the predicted hardening modulus for the elongated grains is lower than the more conventional microstructure.It is difﬁcult to draw quantitative conclusions on the effect that the grain boundary has on the stress-strain behavior. However, it can be said that the grain shape and size both have a discernible effect on the mechanical behavior of the PAM part. It is also known that grain size and shape are inﬂuenced by the process parameters involved in producing the PAM part.

It is also known that grain size and shape are inﬂuenced by the process parameters involved in producing the PAM part. The combination of these facts leads to the conclusion that the process parameters can be directly related to the mechanical properties of the PAM produced part.ConclusionsThe work in the present paper has described an overview of the implementation of a constitutive model that incorporates grain size and shape effects. Synthetic microsturcture repre- sentative volume elements were generated using a continuum diffuse interface model. RVEs of three different aspect ratios, representative of PAM microstructures, were generated and then con- verted to ﬁnite element meshes. The FE models were implemented in Abaqus/Standard with the crystal plasticity model implemented in a UMAT.

The FE models were implemented in Abaqus/Standard with the crystal plasticity model implemented in a UMAT. The FE models of the RVEs were used to com- plete a parametric study to determine the effects of grain size and shape on effective mechanical properties of the RVE.This material is declared a work of the U.S. Government and is not subject to copyright protection 1758 in the United States. Approved for public release; distribution is unlimited.κ0 = 1, Yκ0 = 3, Yκ0 = 5.5, Yκ0 = 1, Zκ0 = 3, Zκ0 = 5.5, Zκ0 = 1, X-Z.κ0 = 3, X-Z.κ0 = 5.5, X-Z.κ0 = 1, Y-Z.κ0 = 3, Y-Z.κ0 = 5.5, Y-Z.Figure 5: Results of the parametric study varying RVE anisotropy factor (κ0), grain boundary effect applied, and load direction.This material is declared a work of the U.S. Government and is not subject to copyright protection 1759 in the United States.

Government and is not subject to copyright protection 1759 in the United States. Approved for public release; distribution is unlimited.As part of the parametric study, three synthetic RVEs were simulated under four loading conditions with and without grain boundary effects applied. The results showed that when both the size and shape effects are considered, the loading in the direction of the elongated grains produced superior strength compared to the out-of-plane loading.

The results showed that when both the size and shape effects are considered, the loading in the direction of the elongated grains produced superior strength compared to the out-of-plane loading. Finally, it was concluded that the grain boundary effect had the most signiﬁcant effect immediately post yield with the effect diminishing as the material strength reached the saturated strength.With a known connection between process parameters and microstructure, it may be possible to utilize the framework and the insight gained from this work to optimize a microstructure so that it would exhibit superior yield strength while maintaining a higher hardening modulus compared to traditionally manufactured microstructures. The optimized structure could then be related back to speciﬁc PAM process parameters and the parts exhibiting the superior behavior could be built.AcknowledgmentsWe acknowledge the Sabbatical Faculty Fellowship with the Technology Management Training Group, Inc.

(TMT Group), and support by US Naval Research Laboratory and Clarkson University for AA. Partial support for this project was provided by the Ofﬁce of Naval Research (ONR) through the Naval Research Laboratorys Basic Research Program. This work was supported in part by a grant of computer time from the DOD High Performance Computing Modernization Program at the Army Engineer Research and Development Center (ERDC) DoD Supercomputing Resource Centers (DSRC).References[1] J.-P. Kruth, L. Froyen, J. Van Vaerenbergh, P. Mercelis, M. Rombouts, and B. Lauwers, “Selective laser melting of iron-based powder,” Journal of Materials Processing Technology, vol. 149, no. 1, pp. 616–622, 2004.[2] G. Kasperovich and J. Hausmann, “Improvement of fatigue resistance and ductility of tial6v4 processed by selective laser melting,” Journal of Materials Processing Technology, vol. 220, pp. 202 – 214, 2015.[3] R. Li, Y. Shi, Z. Wang, L. Wang, J. Liu, and W.

220, pp. 202 – 214, 2015.[3] R. Li, Y. Shi, Z. Wang, L. Wang, J. Liu, and W. Jiang, “Densiﬁcation behavior of gas and water atomized 316l stainless steel powder during selective laser melting,” Applied Surface Science, vol. 256, no. 13, pp. 4350 – 4356, 2010.[4] M. Matsumoto, M. Shiomi, K. Osakada, and F. Abe, “Finite element analysis of single layer forming on metallic powder bed in rapid prototyping by selective laser processing,” International Journal of Machine Tools and Manufacture, 2002.This material is declared a work of the U.S. Government and is not subject to copyright protection 1760 in the United States. Approved for public release; distribution is unlimited.[5] R. B. Patil and V. Yadava, “Finite element analysis of temperature distribution in single metal- lic powder layer during metal laser sintering,” International Journal of Machine Tools and Manufacture, vol. 47, no. 7, pp. 1069–1080, 2007.[6] D. Zhang, Q. Cai, J. Liu, L. Zhang, and R.

136, no. 3, 2014.[16] A. Achuthan, A. Iliopoulos, J. Michopoulos, R. Saunders, and A. Bagchi, “Towards a con- stitutive model that encapsulates microstructural features induced by powder additive man- ufacturing,” in ASME 2017 International Design Engineering Technical Conferences andThis material is declared a work of the U.S. Government and is not subject to copyright protection 1761 in the United States. Approved for public release; distribution is unlimited.Computers and Information in Engineering Conference, American Society of Mechanical Engineers, 2017.[17] M. G. Moghaddam, A. Achuthan, B. A. Bednarcyk, S. M. Arnold, and E. J. Pineda, “Devel- opment of a precipitate size-dependent crystal plasticity constitutive model for two-phase ma- terials and its implementation on a multi-scale computational framework,” Materials Science and Engineering: A, vol. 651, pp. 893 – 903, 2016.[18] M. Ghorbani Moghaddam, A. Achuthan, B. A. Bednarcyk, S. M. Arnold, and E. J.

30, no. 2, pp. 411 – 417, 1982.This material is declared a work of the U.S. Government and is not subject to copyright protection 1762 in the United States. Approved for public release; distribution is unlimited..

lid Freeform Fabrication 2021: Proceedings of the 32nd Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperTHE INFLUENCE OF GRAIN SIZE DISTRIBUTION OF PA12 ON KEY STEPS OF THE POLYMER LASER SINTERING PROCESS Jens P.W. Sessega,b,*, Paul Riedmanna,c, Sybille Fischera, Hans-Joachim Schmidb a EOS GmbH, Robert-Stirling-Ring 1, 82152 Krailling, Germany b Paderborn Institute for Additive Fabrication, University of Paderborn, Pohlweg 55, 33098 Paderborn, Germany c Department of Mechanical Engineering, Technical University of Munich, Boltzmannstrasse 15, 85748 Garching, GermanyAbstractAs the industry pushes for higher resolution laser sintering, finer and finer powders are required. Yet, this also changes the way powders behave during the process. In this project, the influence of finer particles on the dosing and coating process during laser sintering is being investigated.

Yet, this also changes the way powders behave during the process. In this project, the influence of finer particles on the dosing and coating process during laser sintering is being investigated. PA12 laser sintering powder without flow additives was sieved to four fractions with grain sizes with a d50 of 42, 54, 61 and 66 µm. These fractions were characterized regarding powder flowability by FT4 powder rheometer. The dosing and coating behavior of these fractions was tested in separate experiments on an EOS P395, where the coating experiments were performed at both RT and 175°C. Furthermore, test jobs were built with the different powder fractions.

Furthermore, test jobs were built with the different powder fractions. The larger portion of fines in the finest fraction dominated the powder properties and led to significantly worse powder behavior during the dosing and coating steps, as predicted by analytical measurements.KeywordsAdditive Manufacturing, Laser sintering, Polymer, Grain Size Distribution, Powder Rheology, Polyamide 12IntroductionLaser sintering is an uprising technology with an expanding range of applications, but also a lot of untapped potential. One direction in which it is being improved, is a finer resolution, which at some point inevitably leads to the topic of finer powders. Finer powders have two advantages that can help improve part quality. One advantage is, that it allows for a thinner layer thickness. This directly impacts the resolution in Z-direction, as it can make the stepping effect of sloped surfaces finer. The other one is the positive influence on the surface quality [1].

This directly impacts the resolution in Z-direction, as it can make the stepping effect of sloped surfaces finer. The other one is the positive influence on the surface quality [1]. Unfortunately, a finer particle distribution, among other parameters, also leads to decreased powder flow properties [2-8]. Good powder flowability is essential for laser sintering. Primarily, the dosing and coating steps need to be ensured. If one of these steps fails, no parts are being built or part quality is compromised. Additionally, powders with better flow properties pack more efficiently, which was shown to positively contribute to better mechanical properties [2] and surface quality [9]. Unfortunately, powder flow cannot be described as a single value, as a powder behaves differently under different conditions. Powders can have significantly different behavior depending on the stress state and whether they are evaluated under static or dynamic conditions [3].

Powders can have significantly different behavior depending on the stress state and whether they are evaluated under static or dynamic conditions [3]. Hence, a powder657that can be easily dosed, may show difficulties in the coating process. Yet, both are required in order to create a decent powder bed.The theoretical maximum packing density is independent of a spheres’ diameter and the material. But it can be improved by a mixture of differently sized particles where smaller particles fill the voids created by densely packed larger particles. In theory, a packing density of near 100% could be reached this way [10]. This would require a multimodal powder with particle sizes ranging over many orders of magnitude. Practically, this kind of powder would be quite difficult to produce and would probably have terrible flow properties.

Practically, this kind of powder would be quite difficult to produce and would probably have terrible flow properties. With the currently most common layer thick- nesses of 100 to 120 µm, the upper particle size is roughly limited by that value while the lower limit is roughly given by the particle size, when fine particles make the powder too adhesive, a size ratio of 1:3 is probably the limit of what could be expected as grain size difference. Karapatis et al. [11] tested the increase in layer density for bimodal metal powders with different concentrations. A 1:2 ratio did not result in a significant improvement, while with a 1:5 ratio an increase in layer density from about 55% to 61% was found. They also found that curves for bulk and layer density are similar, indicating that the same processes occurred, although with less efficiency in the layer case. Nevertheless, it should be kept in mind that Karapatis tested metal powders, while this study investigates polymer powders.

Nevertheless, it should be kept in mind that Karapatis tested metal powders, while this study investigates polymer powders. The difference in material selection might cause a difference, as the flowability depends of the ratio between adhesion and gravitational forces. Therefore, the effect is most likely less pronounced for polymer powders.Finer particles are required for high-resolution laser sintering parts. Yet their advantages fall flat if powder flowability degrades to a point where it negatively impacts the process or the powder gets entirely unprocessable. In order to mitigate this, it is necessary to understand during which step of the powder application process it fails and the reason for it, in order to subsequently solve the issue by selectively improving other powder parameters, which is what this work is investigating.ExperimentalFig. 1 a) Powder size distribution of the tested PA12 powder fractions, the blue line representing the d50.

1 a) Powder size distribution of the tested PA12 powder fractions, the blue line representing the d50. Four powders, which were fractionated by sieving, and the original unsieved powder were658tested. b) and c) Schematic of isolated dosing and coating experiments on an EOS P395. d) Schematic representation of build job including tensile bars (green), powder boxes (blue), part density cuboids (purple) and a freeform part (red).MaterialA polyamide 12 (PA12) powder without any added flow agent was used for the followingexperiments.Powder fractionationFour powder fractions with different grain size distributions were created by sieving using an air-jet sieve (SLS 200, Siebtechnik GmbH, Germany). During sieving, static neutralizing bars (MEB, SIMCO B.V. Netherlands) were used to facilitate sieving.

During sieving, static neutralizing bars (MEB, SIMCO B.V. Netherlands) were used to facilitate sieving. Sieves with mesh sizes of 80 μm, 63 μm and 40 μm were used.AnalyticsPowder rheology was determined by measuring the compression, aeration and unconfined yield strength on an FT4 powder rheometer (Freeman technology, Great Britain). The shearing measurements were done according to ASTM D7891-15. Pre-shear normal stress was set to 15kPa. The avalanche angle was measured with a Revolution Powder Analyzer (Mercury Scientific Inc., USA). 250 Avalanches were measured per sample at a rotation speed of 0.5 RPM and an inner drum diameter of 100 mm. Part density of the sintered cuboids was measured according to the Archimedes principle with a 770-60 scale (KERN & SOHN GmbH, Germany) in deionized water with a drop of dishwasher soap to decrease the surface tension of water.

Tensile testing was done at norm conditions and according to DIN EN ISO 527-1:2019 employing a Zwick Z005 (ZwickRoell GmbH & Co. KG, Germany).Dosing experimentTo investigate the dosing behavior of different powder fractions, this step was tested separately on a laser sintering machine (EOS P395, EOS GmbH, Germany). The powder was placed in the dosing bin 16h before the experiment. Fluidization in the dosing bin was set to 10 l/min in pulsed operation mode. 2 kg of powder were put in the dosing bin and conditioned for one hour with the fluidization turned on before starting the measurement. For each data point 10 units were dosed. One dosing unit is defined by the machine and corresponds to a theoretical maximum of 9.9 cm³. The dosed powder was collected and weighed with a scale, model WEJ 3000-2 (KERN & SOHN GmbH, Germany). This experiment was repeated 30 times for each of the 4 powder fractions as well as the unsieved powder.

This experiment was repeated 30 times for each of the 4 powder fractions as well as the unsieved powder. To minimize variation of powder settling due to the pulsed operation of fluidization, dosing was always done 5-10 s after fluidization had finished.Coating experimentPowder coating was tested by manually filling either 12.5g (referred to as “low filling”), 25 g (“normal filling”) or 37.5 g (“high filling”) of powder in the empty recoater, where the amount of 25 g was chosen such, that the recoater filling corresponded to the usual filling in a standard building process. The recoater was moved over the powder bed at a speed of 150 mm/s once and the remainng powder in the recoater was collected and weighed. The difference in weight, was assumed to be the powder deposited. The amount deposited was recorded 5 times for consecutive single layers. When changing recoater filling level, two layers were laid down unrecorded to reach equilibrium.

The amount deposited was recorded 5 times for consecutive single layers. When changing recoater filling level, two layers were laid down unrecorded to reach equilibrium. It was verified afterwards, that an equilibrium had been reached. The measurement659was performed at room temperature, as well as at a temperature of 175°C, which is close to the process temperature for PA12. For the experiments at 175 °C, the powder was preheated to 70°C for 2h in an oven under N2 atmosphere to account for the warming of the powder bin during the build process. The powder movement in the recoater during the coating steps was recorded with a GoPro Hero 7 (GoPro Inc., USA) mounted on the recoater.Job setupThe build job is shown in Fig. 1 d).

The powder movement in the recoater during the coating steps was recorded with a GoPro Hero 7 (GoPro Inc., USA) mounted on the recoater.Job setupThe build job is shown in Fig. 1 d). It includes 10 tensile bars in XY-direction according to DIN EN ISO 20753 type A1, 10 cuboids for determination of part density, 4 hollow boxes to measure the powder bed density and a freeform part for surface evaluation.BuildThe build was performed on an EOS P395 (EOS GmbH, Germany, Software: PSW 3.6.91) laser sintering machine, employing the PA2221_120_000 Default-Job, with the PA12 powder fractions described above and the reference powder. A temperature search determined the optimal build temperature to be 179°C. In all cases a 6 mm powder layer was laid down before preheating the machine for 2 hours prior to the job start. The powder in the dosing bin was fluidized in pulsed operation mode with a gas flow of 10 l/min.

The powder in the dosing bin was fluidized in pulsed operation mode with a gas flow of 10 l/min. The build job was closely monitored and in case of an underfeed, more powder was added manually in order to avoid a build interruption.660Results and discussionFig. 2 a/b) Conditioned bulk density, and compressibility at both 2 and 16 kPa pressure, of the different powder fractions. c/d) Unconfined yield strength and the inversely behaving flow factor measured with a pre-shear normal stress of 15 kPa. e) The avalanche angle with the 95th percentile emphasized. f) Cumulative energy for a preset stirring program over fluidization flow.The avalanche angle was measured at a slower speed than the speed of powder movement one would expect inside the recoater. Yet, with no normal stress on the powder, the powder flow conditions should be close to process behavior.

Yet, with no normal stress on the powder, the powder flow conditions should be close to process behavior. As the avalanche angle is not a specific value, but rather a distribution (which in fact is not an inaccuracy of the measurement but a significant effect of the powder), not the median of each powder, but the 95th percentile was compared. It describes the angle of the powder surface at which an avalanche has reliably started. Considering the 99th percentile or largest recorded value would have been too prone to inconsistencies. Powders 2,3 and 4 had slightly increasing avalanche angles with finer powder distributions, but have had most avalanches by 60°, while powder fraction 1 had the 95th percentile at 71° with the last avalanches only breaking at 80°.

Rheological behavior of the unsieved powder in all measurements is very similar to the powder fraction 2 which has a similar median value, but slightly narrower grain size distribution.Contrary to the other measurements that had better flowability with larger particles, aeration of powders with finer particles was easier facilitated by the larger surface to mass ratio. Powder fraction 1 with the largest cohesion was easiest to fluidize and the only powder to reach an equilibrium within the range of the measurement. Weirdly, stirring energy for this powder was also661smallest with no air speed, despite its highest cohesion. This effect could be caused by the inferior powder packing behavior of this powder fraction.DosingDosing of the different powder fractions showed three effects (Fig. 3a). The amount of powder dosed decreased with increasing iterations for all powders.

3a). The amount of powder dosed decreased with increasing iterations for all powders. This can be attributed to the observed powder build-up on the dosing rolls which effectively reduced the volume each dosing unit can convey. Only powder fraction 1 reached an equilibrium during the experiment, yet at a pretty low level of around 10 g/iteration. The increased caking behavior of fraction 1 is in accordance with the higher compressibility and larger unconfined yield strength. This means, that under pressure the powder tends to rearrange more and the resulting powder cake is more stable.Fractions 2-4 and the unsieved fraction showed similar behavior to each other during the whole experiment.

The finest powder fraction 1 started at a lower level and quickly reached an equilibrium with significant fluctuation at roughly 10 g/10 dosing units, which is roughly a third of what the other powders started with and only around 25% of the theoretical maximum of roughly 40 g/10 dosing units (with a bulk density of PA12 of 0,41 g/cm³ (Fig. 2b)).Two of the powder fractions (1 and 2) showed a dip in dosing at step 5 and 6. These were caused by bridging of the powder which had compacted over time during the resting phase before the experiment. The powder in the parallel gap between the bridge and the powder roll could still be dosed, before the powder bridge became problematic. Bridging could be solved by hitting the dosing bins, visibly collapsing the powder bridge. The reason, that the dip of powder fraction 1 did not go to near-zero was a result of the powder bin being hit when the effect was clearly visible, breaking powder bridges and, hence continuing the experiment.

No significant caking or powder build-up was observed during later build jobs with any of the powders. This was probably prevented by the increased temperature and dry conditions during the building process.662Fig. 3 a) Stability of the amount of powder dosed during subsequent steps of 10 dosing units, depending on the grain size distribution b) The amount of powder deposited on the powder bed depending on the grain size distribution and filling level of the recoater. This experiment was performed at room temperature (RT) as well as at operating temperature of 175 °C (HT). c) Classification of different powder movement behaviors in the recoater. 0: No visible powder movement. 1: Only slight scooping at the blade. 1*: No circulation, but rolling of agglomerates leading to visible vibrations. 2: Scooping motion without breaking. 3: As 2 but with chunks breaking off. 4: Circulation with chunks falling off regularly. 5: Near-continuous powder movement with small avalanches.

2: Scooping motion without breaking. 3: As 2 but with chunks breaking off. 4: Circulation with chunks falling off regularly. 5: Near-continuous powder movement with small avalanches. d) Classification of the powder movement according to c) for the different experiments at room temperature (RT) and elevated temperature (HT) with a variation in recoater filling rate, dependent on the powder fraction.Fig. 4 Powder bed quality produced from the powder fractions at 175°C with 25g powder in recoater.663CoatingThe amount of powder deposited in one coating step was measured for different powder fractions and recoater filling levels (Fig. 3b). This experiment was done at room temperature and at process temperature. In general, less powder was deposited at process temperature than at room temperature. Moreover, at room temperature, the amount of powder deposited was largely the same for all powder fractions, including the notoriously static fraction 1.

Moreover, at room temperature, the amount of powder deposited was largely the same for all powder fractions, including the notoriously static fraction 1. Yet, a fuller recoater deposited slightly more powder. This was different, though, when repeating the experiment at operating temperature. Still, a fuller recoater had a positive influence on the amount of powder deposited, but additionally there was a difference between the five powder fractions. While of the powder fractions 2-4, which already showed very similar avalanche angles and similar unconfined yield strengths, also similar amounts of powder were deposited, powder fraction 1 performed significantly worse and only a third of the powder compared to the other fractions was deposited in the case of a normal recoater filling level of 25 g. A high filling level could partially mitigate this problem for this powder fraction.

A high filling level could partially mitigate this problem for this powder fraction. The unsieved powder behaved similar to the coarser powder fractions, except for the high filling level at operating temperatures which had two outliers with a large deposition, facilitated by an uneven powder bed.Of special interest is the comparison of the amount of powder deposited in the different cases and the powder movement in the recoater (Fig. 3c/d). At room temperature all powders showed powder movement in the recoater. This was mostly independent of the filling level, except for fraction 1. This fraction showed less movement with a fuller recoater with only vibrations visible on the surface in the worst case and agglomerates being rolled between the powder and the powder bed. At operating temperatures, powder movement was generally reduced.

At operating temperatures, powder movement was generally reduced. The coarser powder fractions still showed adequate movement, but the finest fraction showed no powder movement anymore, rather agglomerates rolled as whole entities. As the powder mainly consisted of larger agglomerates, a thin layer of powder could not be coated. It is assumed that instead powder was scraped off of the agglomerates where it touched the powder bed, resulting in some deposition. Meanwhile slight irregularities, where the agglomerates did not contact the powder bed, resulted in no powder being deposited at all, which added up to grooves in the powder bed that got deeper with every layer (Fig. 4). The high recoater filling level of powder fraction 1 could partially mitigate this problem by more pressure on the powder bed under the powder’s own weight. The behavior of the unsieved reference powder was somewhere in between the coarse and the fine fractions.

The behavior of the unsieved reference powder was somewhere in between the coarse and the fine fractions. In the case of a low filling level the powder seemed to float over the powder bed with no interaction at all, and only with more weight pushing down the powder a movement could be introduced. Overall, powders that showed powder movement of type 2 or lower resulted in a textured powder bed. More powder movement in the recoater led to a good powder bed (Fig. 4).BuildThe reference powder as well as the three coarser powders were adequately processable. While in single coating tests, coating of powder fraction 1 was successful and data could be obtained, process behavior was by far insufficient for a build job. The reduced powder dosage resulted in a constant underfeed which had to be counteracted by irregular extra dosing units. The bad coating behavior at operating temperatures then led to grooves in the powder bed in which the exposed parts would not result in a coherent part.

The bad coating behavior at operating temperatures then led to grooves in the powder bed in which the exposed parts would not result in a coherent part. These reasons quickly resulted in an entire job crash. Hence, unfortunately no mechanical properties could be determined of this powder. This was already indicated by analytical measurements, which showed a significantly inferior rheological behavior which was clearly visible in the dosing and coating experiments.664The powder bed density, as seen in Fig. 5a), shows the same trend as the conditioned bulk density, where the three powder fractions had a similar density. Meanwhile the unsieved powder showed a higher powder bed density, though this effect was stronger than the conditioned bulk density suggested.

Meanwhile the unsieved powder showed a higher powder bed density, though this effect was stronger than the conditioned bulk density suggested. Part density of the cuboids and Young’s modulus both showed a slightly increasing trend with coarser powder fractions and the unsieved powder being roughly in the middle, and for the Young’s modulus with a larger standard deviation. The difference between the trends of powder bed density and part density, could be explained by the fact that in a powder bed the grains need to pack with solid particles, hence the good correlation with the conditioned bulk density measurement. Yet, in a part the new powder is being pressed into the underlying still molten layer. The differences between the part density and Young’s modulus are minor though, but in the case of the part density still larger than the standard deviation.

The differences between the part density and Young’s modulus are minor though, but in the case of the part density still larger than the standard deviation. They could show a tendency of part density and Modulus being influenced by the increasing bulk density of the fractions but could also be caused by variations in-between different builds. Elongation at break was not significantly different for the powders that were processable.Fig. 5 Data gathered from the test parts of the build jobs. Data points for fraction 1 have been left empty as the fraction could not be processed. a) Powder bed density in measured from powder enclosed in hollow parts. b) Part density of test parts. c/d) Mechanical properties of tensile bars.It is unfortunate, that the powder fraction 1 performed so much worse than the other powder fractions and could not be processed.

c/d) Mechanical properties of tensile bars.It is unfortunate, that the powder fraction 1 performed so much worse than the other powder fractions and could not be processed. Nevertheless, the builds jobs indicate, that the better flowability of coarser powder fractions result in slightly improved part density and stiffness. As the experiments were not repeated, this effect could be coincidental. More importantly, it could be shown that the fines define the powder behavior significantly more than the average grain size. On665the positive side, this does indicate, that finer powder fractions, than are currently common, can still yield promising results, as long as the finest fraction is removed, especially when flow additives are added to the powders. The presented experiments were performed without any flow agents, which are commonly used in laser sintering materials. This means, that powders are expected to perform significantly better when flow agent is being added.

This means, that powders are expected to perform significantly better when flow agent is being added. Moreover, it can be assumed, that it is viable for powders with finer grain size distribution to perform well for laser sintering and other powder bed fusion processes.Conclusion and OutlookThis work showed the influence of particle size distribution on powder rheology through different steps of the powder application process on a laser sintering machine as well as the influence on properties of test specimens built. A standard PA12 powder was sieved into 4 fractions with different particle size distributions. It could be shown that the bulk density of powder fractions with a narrower particle size distribution was slightly lower than for the unsieved powder with a wider particle size distribution. This can be explained by a better packing due to filling up gaps with smaller particles.

This can be explained by a better packing due to filling up gaps with smaller particles. The results from fraction 1 (<40 µm) showed impressively the influence of the finest fraction of particles. Here, obviously intermolecular interactions and attraction dominate the powder rheology and lead to a powder which behaves sticky and tends to build agglomerates.The differences in powder rheology were transferable to the consecutive process steps of powder application in the laser sintering machine, where the fine fraction had the highest tendency for bridging in the dosing system and showed little movement in the recoater. The dosing behavior was in line with the results from the unconfined yield strength and compressibility measurements.

The dosing behavior was in line with the results from the unconfined yield strength and compressibility measurements. Comparing the coating process with the resulting properties of the test specimens indicated that the movement in the recoater is critical for the creation of a smooth and dense powder bed which again is inevitable for dense parts with decent mechanical properties. The powder movement inside the recoating system correlated well with the analytical measures determined from shear-cell, powder rheometer and the Revolution Powder Analyzer, which therefore can be employed to improve the prediction of powder processability based on powder analytics.AcknowledgementsThe authors acknowledge Prof. Katrin Wudy (Professorship of Laser-based Additive Manufacturing; TUM) for supervising Mr. Riedmann’s project and the insightful discussions.666References[1] P. B. Bacchewar, S. K. Singhal and P. M.

10, no. 2, pp. 203- 208, 2012.[8] H. Shi, R. Mohanty, S. Chakravarty, R. Cabiscol, M. Morgeneyer, H. Zetzener, J. Y. Ooi, A. Kwade, S. Luding and V. Magnanimo, "Effect of Particle Size and Cohesion on Powder Yielding and Flow," KONA Powder and Particle Journal, pp. 226-250, 2018.[9] L. Meyer, A. Wegner and G. Witt, "Influence of the Ratio between the Translation and Contra-Rotating Coating Mechanism on different Laser Sintering Materials and their Packing Density," in Solid Freeform Fabrication 2017, Austin, 2017.[10] A. Averardi, C. Cola, S. E. Zeltmann and N. Gupta, "Effects of particle size distribution on the packing of powder beds: A critical discussion relevant to additive manufacturing," Materials Today Communications, vol. 24, no. 2352-4928, p. 100964, 2020.667[11] N. P. Karapatis, G. Egger, P. Gygax and R. Glardon, "Optimization of powder layer density in selective laser sintering," in International Solid Freeform Fabrication Symposium, Austin, 1999.[12] A. W.

P. Karapatis, G. Egger, P. Gygax and R. Glardon, "Optimization of powder layer density in selective laser sintering," in International Solid Freeform Fabrication Symposium, Austin, 1999.[12] A. W. Jenike, "Storage and flow of solids," Bulletin No. 123 of the Utah EngineeringExperiment Station, vol. 53, no. 26, 1976.668.

fluence of heat-treatment on Selective Laser Melting products – e.g. Ti6Al4VM. Thöne1,2, S. Leuders1,2, A. Riemer3, T. Tröster2, H.A. Richard31 Direct Manufacturing Research Center DMRC, Mersinweg 3, 33098 Paderborn, Germany 2 Automotive Lightweight Constructions LiA, Pohlweg 47-49, 33098 Paderborn, Germany 3 Institute of Applied Mechanics FAM, Pohlweg 47-49, 33098 Paderborn, GermanyAbstractUsually additive manufactured metal parts are showing a different mechanical behavior compared to conventionally produced parts used the same material. Apart from process-related macroscopic part imperfections (pores, surface roughness, etc.) the microstructure has a decisive influence on the mechanical properties of the materials. Thus, in order to optimize mechanical properties of metal parts a heat treatment for changing microstructures is routinely applied in most production lines to meet the product requirements.

Thus, in order to optimize mechanical properties of metal parts a heat treatment for changing microstructures is routinely applied in most production lines to meet the product requirements. By means of the Titanium alloy Ti6Al4V the optimization of the static- and the fracture mechanical behavior by changing the microstructure with a heat treatment after the SLM process is discussed on the present work.IntroductionThe additive manufacturing (AM) details for metals has been published for about 20 years. First additive manufacturing processes were used primarily for rapid prototyping in variants of polymers and some metals [1]. Today, most of industry indicated their interest at manufacturing production by using selective laser melting (SLM) system work, rather than using AM for rapid prototyping. [2].

Today, most of industry indicated their interest at manufacturing production by using selective laser melting (SLM) system work, rather than using AM for rapid prototyping. [2]. The mechanical behavior of SLM parts becomes vital information necessary to be used for component production.It is important to accept the fact that the mechanical behavior of SLM parts are different to conventionally produced parts. Selective Laser Melting is a laser build-up welding technology with high cooling rates that result in a metal microstructure which is not comparable to metal casting or forged parts [7]. The dimension of difference from cast or forged parts varies depending upon the material being used. In all cases, the built SLM parts possess internal stresses, where the difference in mechanical behavior can be either an advantage or a drawback.In most cases it is desirable to produce parts without internal stress.

Currently, SLM parts employ a heat treatment for reduction of internal stresses but additional advantages are also possible from the heat treatment. The example of SLM parts out of Ti6Al4V will illustrate the mechanical property possibilities resulting from heat treatments. The primary aim of a heat treatment in this case of Ti6Al4V is the reduction of internal stresses. Due to the mechanical behavior of Ti6Al4V as build SLM parts, the heat treatment also seems to be essential to increase the mechanical behavior, e.g. the breaking elongation, the fatigue performance, etc.Experimental DetailsThe investigation of material properties resulting from nine different types of heat treatments of Ti6Al4V SLM parts is presented within this body of work. The Ti6Al4V experimental tensile test bars and cantilever test specimens are built on a SLM250HL system from SLM Solutions GmbH in Germany with a 400W Ytterbium fiber laser. The Ti6Al4V powder particle size has an average value of 40µm.

The Ti6Al4V powder particle size has an average value of 40µm. All used specimens were built using a 175W Laser Power setting and 30µm layer thickness. That is standard settings for laser melting of Ti6Al4V supported by SLM Solution GmbH.The parameters for all nine tested heat treatments are shown in Table 1. The first variants of the heat treatments are performed at 750°C up to 850°C for 2h under an Argon gas atmosphere and cooled down inside the furnace. The next variants of the heat treatments are performed in vacuum with 950°C and 1050°C for 2h cooled down inside the furnace. Vacuum is used because the DMRC equipment is unsuitable for using Argon492gas above a temperature of 850°C. Comparison tests between vacuum and Argon treated specimens show no important deviations.

Comparison tests between vacuum and Argon treated specimens show no important deviations. The last heat treatment variant was a hot isostatic pressing (HIP) process at 920°C for 2h.Table 1: Parameters of heat treatment.Temperature / °C Time / h Atmosphere1 20 - -2 750 23 775 24 800 25 825 26 850 27 950 28 1050 2Argon Argon Argon Argon Argon Vacuum Vacuum Argon9 HIP 2To illustrate the influence of heat treatment on SLM-Ti64-parts five experimental tests will be performed and explained within this paper; testing of internal stress with the help of cantilevers, testing of mechanical behavior with quasi-static tensile tests, HCF-tests or fracture tests and micrograph analysis.The internal stress of SLM-parts cannot be measured at the DMRC so the deformation by cutting the cantilevers from the building platform is measured. The four cantilevers have a length of 44 mm and a thickness of 4mm (Figure 1).

The four cantilevers have a length of 44 mm and a thickness of 4mm (Figure 1). They are heated up in 25°C steps from 750°C to 850°C together with the building platform. A comparison with a non-heat-treated cantilever provides some indication of the development of internal stresses.All quasi-static tensile tests are performed on a Universal Testing Machine Instron 5569 according to the ISO 6892-1:2009 standard [3]. The specimens (Form A20x4) are built with the gauge length in the z-direction and tested as build deviating from standard. A minimum of five specimens will be tested for each level of heat treatment temperature. The load will be measured with a 50 kN load cell and the elongation with an optical extensometer and with a measuring length of 20mm conducted at room temperature.

The load will be measured with a 50 kN load cell and the elongation with an optical extensometer and with a measuring length of 20mm conducted at room temperature. The tensile stress Rm and the breaking elongation At are evaluated and presented within this paper.A characterization of the microstructure of both the SLM-processed and the heat treated specimens provide information about microstructural behavior resulting from the heat treatment. A connection with the mechanical behavior and the microstructure is detected by an electron backscatter diffraction (EBSD) system. The EBSD scans are conducted at 20 kV on electro polished specimens with a scanning area of 30x30 µm with a step size of 0.15 µm. The scans will be differentiating between hexagonal grid structure (α’,α) and body centered grid structure (β).In addition to the quasi-static tensile tests, high cycle fatigue tests (HCF) are performed.

The HCF-tests are run under the same atmosphere conditions as the quasi-static tests, at room temperature, and are performed on a Zwick/Roell Amsler HB250fatigue strength testing machine. The test runs are force-controlled with amplitude of 600 MPa and at a cylce frequency of 40 Hz. The mean force of the force cycle waveform is 0 N with a reversed force ratio R=-1. The round specimens are built in the z-direction and based on the standard ASTM E466 – 07.The analyses of the crack growth behavior and fracture tests are run on an Instron testing machine ElectroPlusTM under room temperature (20°C). The test specimens are manufactured according to the standard of ASTM E647-08 and were built in the z-direction.

The test specimens are manufactured according to the standard of ASTM E647-08 and were built in the z-direction. The tests are run with a force ratio of R=0.1 and in the area of K-decreasing (area of crack initiation) with lower crack growth rates by a frequency of 40 Hz and in the area with higher crack growth rates by a frequency of 10 Hz.ResultsThe first aim of heat treatment is for the reduction of internal stresses. Strain measurements after cutting the support structure of the cantilevers are plotted in Figure 1. Figure 1 shows a comparison of the non-heat- treated specimen (20°C) where there is no considerable bending measured. The relation between internal stresses and the bending of cantilevers is a good index for measuring the reduction of internal stress.

The relation between internal stresses and the bending of cantilevers is a good index for measuring the reduction of internal stress. For all the tested temperature levels from 750°C to 850°C, the measurable strain reduces from 0.32mm to 0.05mm or less.4930,350,30,25]m m0,2[ n i a r tS0,1520mm0,10,05020°C750°C775°C800°C Heat treament825°C850°CFigure 1: Bending of Ti64 cantilevers after heat treatment.The results of tensile tests are shown in Figure 2 and Figure 3. All results are produced with specimens that were tested as built in the z-direction and compared non-heat-treated specimens. Figure 2 shows the measured tensile strength Rm and Figure 3 shows the appropriate breaking elongation At. The first bar in both diagrams representing the specimen without any post process shows a nearly brittle structure.

The first bar in both diagrams representing the specimen without any post process shows a nearly brittle structure. Due to that, while the decreasing of tensile strength from 1080 MPa to 945 MPa seems to be fractional (12.5%) the increasing of breaking elongation from 1.6% to 11.6% (625%) is reached. For example international standards for implants for surgery are claiming special minimum values for the tensile strength and the breaking elongation especially for Ti6Al4V [4]. Therefor the tensile strength should realize a minimum of 860 MPa and the breaking elongation a minimum of 10%. In case of SLM produce parts this requirements are fulfilled by heat-treatment.The reason of such an increasing of breaking elongation can’t be described with the reduction of internal stress shown in Figure 1. Internal stress isn’t already detectable after a 750°C heat treatment but also with an 850°C heat-treatment a maximum of only 5.2% breaking elongation in average can be reached.

Internal stress isn’t already detectable after a 750°C heat treatment but also with an 850°C heat-treatment a maximum of only 5.2% breaking elongation in average can be reached. Due to this, other effects have to explain the optimization of breaking elongation up to 11.6% in average and a view to micrographs seems to be necessary.494] a P M[s s e r t se l i s n e T]% [ n o i t a g n o l E g n i k a e r b120010008006004002000141210864201.0801.0621.0401.00997220°C750°C800°C850°C950°CFigure 2: Tensile stress of heat-treated specimens10,15,15,23,71,620°C750°C800°C850°C950°CFigure 3: Breaking elongation of heat-treated specimens4959451050°C11,61050°CIn addition to tensile testing, looking at the microstructure to determine the phase development of α and β titanium is also investigated.

A SEM is used to detect structural elements and identify them as α and β constellations with the help of an electron back scatter diffraction (EBSD) scan. Figure 4 shows two EBSD micrographs of a Ti6Al4V specimen as build (left) is compared with a 950°C heat treated specimen. First notice that the laminar structures of the non-heat-treated specimen (left side) are very fine when compared to the 950°C heat-treated specimen (right side). That is one reason for an increasing of breaking elongation because sliding effects are mainly detected between grains [5]. The second effect can be illustrated with the help of EBSD scan and the following determination of hexagonal grid structures (α’, α) and body centered grid structures (β). The difference between the martensitic α’-phase and the α-phase is only a minimal structure distortion and is not detectable with EBSD.

The difference between the martensitic α’-phase and the α-phase is only a minimal structure distortion and is not detectable with EBSD. Due to the high cooling rate inside the SLM process the probability of formation of an α’-phase in the non-heat-treated specimen is very high [6,7,8]. The α-phase and the α’-phase are colored in red where the body centered β-phase is colored in green which were only detected in the heat- treated specimens. The β-phase is mainly detected on the edge of laminar structure and also supported a more ductility behavior of the material [5].Figure 4: Comparison of EBSD scans (30µm x 30µm) between non-heated (left) and after the 950°C heat treated specimens (right)[5].In addition to the analysis of quasi-static behavior it is of interest to take a look at high cycle fatigue behavior (HCF). Table 2 list the results of HCF tests described in the experimental details.

Table 2 list the results of HCF tests described in the experimental details. Table 2 shows the different heat treatments, the frequency of cycles, the stress level and the endurance of fatigue life Nf, together with the standard deviation. For all treatments a minimum of 5 specimens are tested [9]. .Table 2: Results of high cycle fatigue testsTreatment Frequency/Hz Stress/MPa 600 20°C 600 800°C 600 1050°C 600 HIP40 40 40 40Cycles28,900 93,000 290,000 >2,000,000STDV ±11,000 ±16,800 ±76,000496The first three heat-treatments are shown before by the quasi-static tests (Figure 2 and Figure 3). The number of cycles to failure increased with the increasing of temperature heat-treatment. That improvement is similar to what was observed in the quasi-static test results. By changing the microstructure 290,000 cycles are reached. But in no single case it will be reached a level of 2,000,000 cycles to failure.

By changing the microstructure 290,000 cycles are reached. But in no single case it will be reached a level of 2,000,000 cycles to failure. Pores are mostly the initiators of crack growth and result finally in failure of complied parts. Especially pores close to the edge proves to be bad because these pores affect the stress distribution on a micro level [10]. Besides the changing of microstructure it seem to be very important to minimize the number of pores especially at areas close to the edge. To close most of the pores in a post-treatment process in the last case of Table 2 hot isostatic pressing is used.

To close most of the pores in a post-treatment process in the last case of Table 2 hot isostatic pressing is used. A combination between a heat treatment for changing the microstructure and a pressure treatment for closing pores.Also for crack growth curves like shown in Figure 5 with the combination of a heat treatment and apressure treatment, the same characteristics of the parent material from Ti6Al4V can be reached [9].Figure 5: Position of crack growth curves non-treated and heat-treated compared with parent material [9]ConclusionsA close investigation of the mechanical behavior of Ti6Al4V and an evaluation on the influence of heat treatment of SLM-products is given. Due to changes of microstructures three cases of treatment have been identified (Figure 6). In most cases the SLM parts as built possess internal stresses and have a very small sized microstructure.

Due to changes of microstructures three cases of treatment have been identified (Figure 6). In most cases the SLM parts as built possess internal stresses and have a very small sized microstructure. In case 1 (Figure 6) using a heat-treatment is used in the reduction of internal stress without any changing of microstructure. By reduction of internal stress a first beneficial effect is detectable regarding to the mechanical behavior e.g. the increase of elongation at break. By decreasing the internal stress for conservation the shape of SLM parts during cutting parts form the build platform is also a main topic of additive manufacturing. This case 1 of heat treatment can be mostly realized with a short heat-treatment on a low temperature level. Case 2 in Figure 6 is the changing of microstructure of SLM parts by using a higher temperature level or a longer annealing time for heat treatment.

Case 2 in Figure 6 is the changing of microstructure of SLM parts by using a higher temperature level or a longer annealing time for heat treatment. The meaning of changing the microstructure is to enlarge the grain size or the laminar structure (e.g. Ti64) and to change the constellation of microstructure, e.g. from a α’-constellation to a αβ-constellation of Ti6Al4V. And case 3 in Figure 6 is the changing of microstructure combining with reduction of pores by using hot isostatic pressing. In addition to showed optimization of quasi-static behavior in this case also the fatigue properties optimized by closing especially surface near pores.497Figure 6: Three cases of microstructure manipulationAcknowledgmentsThanks for supporting to the DMRC Partners, the University of Paderborn in Germany and theGovernment of North Rhine Westfalia/ GermanyReferences[1] A. Gebhardt, “Generative Fertigungsverfahren” Hanser, 2007, S146[2] J. Gausemeier, N. Echterhoff, M. Kokoschka, M.

Laser Sintering is an additive manufacturing method that offers many advantages over conventional manufacturing techniques of porous structures with well-defined architectures, controllable pore sizes, excellent reproducibility, higher pore interconnectivities and improved mechanical properties can be produced accurately and rapidly.This study describes the morphological and mechanical characterisations of porous Ultra- High Molecular Weight Polyethylene (UHMWPE) laser sintered parts to gain an insight into the correlation of process parameters and the morphological properties of these parts. Laser power was investigated to control the mechanical properties and porosity of the structures.The fabricated parts were characterised through porosity measurements, three point flexural test and scanning electron microscopy (SEM).

X-ray micro-computed tomography (micro-CT) was considered to evaluate the mean internal porosity as well as the size and spatial distribution of pores inside the structure of the UHMWPE parts aiming at a better understanding of the three- dimensional internal morphology of UHMWPE laser-sintered parts. The porosity was then compared with the porosity measured using the helium gas pycnometer method.The results showed a high level of porosity in the UHMWPE laser-sintered parts with a range of 60-65% measured by micro-CT technique and helium gas pycnometer method respectively. There are no significant differences in the results obtained from both techniques and both results fit very well with each other. The results show that flexural strength decreases with an increase in porosity of the sintered parts.Keywords Additive manufacturing, UHMWPE, Laser Sintering, Laser power, Micro computed tomography, Micro-CT8141.

IntroductionAdditive manufacturing (AM) offers remarkable opportunities for producing parts with complex internal structures that are impossible to produce with other manufacturing methods. Therefore, many AM parts have internal geometries that allow optimising the weight, shape and strength of these parts [2]. Laser Sintering (LS) is one of the AM techniques that has the capability to create a controlled complex geometry, both internally and externally, with a good dimensional accuracy and mechanical properties [3], [4]. However, LS parts can be prone to internal porosity. The mechanical properties of porous structures depend not only on the composition and the microstructure of the raw material, but also on the geometrical and morphological properties of the basic pore architecture [5].Ultra High Molecular Weight Polyethylene (UHMWPE) is a thermoplastic polymer which is widely used in medical, industrial and consumer applications [6].

However, UHMWPE cannot be processed easily by injection moulding or melt extrusion due to its very high melt viscosity [7]. UHMWPE can be conventionally processed by compression moulding or ram extrusion [8] resulting in products with a density of 0.939 g/cm3, flexural strength of 14.4 MPa and flexural modulus of 713 MPa [9].Attempts have been made to produce UHMWPE parts using laser sintering with limited success. Rimell and Marquis attempted to laser sinter UHMWPE (GUR 4120 and GUR 4170) using a non-commercial LS machine. The study concluded that multiple layers cannot be produced due to material shrinkage, curling and degradation [10]. Goodridge et al attempted to laser sinter UHMWPE (GUR 4170) using the commercial Vanguard Laser Sintering machine (3D Systems). The mechanical properties of the laser sintered parts were investigated using tensile test and three point bend test.

The mechanical properties of the laser sintered parts were investigated using tensile test and three point bend test. The average Ultimate Tensile Strength and Young’s modulus reported were just above 0.2 MPa and 1.5 GPa respectively and the produced parts have an average flexural strength of 0.52 MPa with a modulus of 18.67 MPa. However, a limited success has been achieved as the sintered parts were only produced by using a precise combination of processing conditions [4].To the authors’ knowledge, no investigation has been reported related to the effect of process parameters on the porosity of UHMWPE laser sintered parts. This study describes the morphological and mechanical characterisations of porous UHMWPE laser sintered parts, built under different conditions, to gain an insight into the correlation of process parameters and the morphology properties of these parts. The laser power was investigated to control the mechanical properties and porosity of the structures.

The laser power was investigated to control the mechanical properties and porosity of the structures. The difference in the porosity levels and the pores morphology between different LS parts were reported. The relationships between laser power, porosity and mechanical properties were discussed.Micro-computed tomography (micro-CT) technique is a promising quality control method in AM that is able to perform non-destructive dimensional measurement of internal structures and non-destructive porosity measurement, which are serious issues in AM [2]. Previous studies investigated the pore size and the porosity distribution using micro-CT. Dupin et al [11] used micro-CT in their work and reported that the laser power has a significant impact on the porosity of polyamide 12 sintered parts.

Dupin et al [11] used micro-CT in their work and reported that the laser power has a significant impact on the porosity of polyamide 12 sintered parts. Rouholamin and Hopkinson [12] stated that micro-CT analysis can815be used as a non-destructive technique to assess the morphology of nylon 12 and measure the porosity and pore size of the laser sintered parts.In this study, x-ray micro-CT was used to evaluate the porosity as well as the size and spatial distribution of pores inside the structure. This research establishes relationships between manufacturing process parameters (i.e. laser power), porosity and with mechanical properties (i.e. flexural properties).2. Materials characterisation and parts manufacturing2.1.

laser power), porosity and with mechanical properties (i.e. flexural properties).2. Materials characterisation and parts manufacturing2.1. Ultra-High Molecular Weight PolyethyleneThe UHMWPE used in this study was supplied from Ticona/Celanese (Germany) as a powder, type GUR 2122, with a reported average particle size of 130 ±20 μm and molecular weight of 4.5 × 106 g/mol with a bulk density of the powder of 0.20 – 0.25 g/cm3 (Celanese GUR® 2122 UHMWPE datasheet). However, the average particle size was determined to be 125 μm (d50), with a size distribution ranges from 46.2 μm (d10) to 293 μm (d90), as measured by laser light scattering using a Mastersizer 3000 (Malvern Instruments, UK) and the dry sample dispersion technique.The peak temperature of the melting and recrystallisation of the UHMWPE powder was determined by Differential Scanning Calorimetry (DSC) (PerkinElmer Pyris DSC 8500) at a heating rate of 10°C/min.

The peak temperatures of the melt and recrystallisation were 141 °C and 117 °C respectively.Scanning electron microscopy (Philips XL-20, Holland, at 15 kV) was used to observe the UHMWPE powder microstructure. The images in Figure 1 show that the particles are non-spherical in shape with a highly agglomerated structure of smaller particles showing a cauliflower like appearance.816Figure 1: SEM micrographs of un-sintered UHMWPE powder with magnifications: a) 1000x, b) 1500x and c) 3000x.2.2. Part manufactureRectangular parts, 80 mm long, 10 mm wide and 4 mm thick, were fabricated using a commercial laser sintering system (EOS P100, Germany). The test parts were oriented in the x-y direction with the long axis parallel to the x-axis (Figure 2).Figure 2: Laser sintered parts of UHMWPE817Four sets of test samples, four samples in each set, were fabricated at different laser powers of 6, 8, 10 and 12 watts.

The laser scan speed, hatch spacing, powder bed temperature, removal chamber temperature and layer thickness were kept constant at 2500 mm/s, 0.15 mm, 142ºC, 135ºC and 0.1 mm respectively, with a double scan count.3. Experimental proceduresThis paper evaluates the successful test parts produced with various laser powers andcharacterises the porosity, mechanical properties, and microstructure of the sintered parts.3.1. Porosity measurementsSmall rectangular specimens with approximately 10 mm length, 8 mm width and 6 mm thickness, were cut out, from the centre and the end, of the laser sintered rectangular parts as shown in Figure 3. Three samples (A, B, and C) were used for density measurements and two samples (A and C) for morphology investigation.

Three samples (A, B, and C) were used for density measurements and two samples (A and C) for morphology investigation. In this study, two methods were used to evaluate the porosity; micro computed tomography (micro-CT) and helium gas pycnometer.Figure 3: Density and morphology samples cut out from the laser sintered parts: Samples A and B from the end and C from the centre3.1.1. Helium gas pycnometerTo determine the part’s porosity, the true density of the UHMWPE powder (ρTrue), the bulk density (ρBulk) and skeletal density (ρSkeletal) of the sintered parts must be known. The bulk density of laser sintered parts was determined using a volumetric method as follows:(1)𝒎𝒎 𝒗𝒗𝐁𝐁𝐁𝐁𝐁𝐁𝐁𝐁𝛒𝛒 The mass of the parts was determined by weighing the samples using a digital balance and the volume of the parts was determined using Vernier callipers.

The true density (or real density) of the UHMWPE powder and the skeletal density of the sintered parts were determined using Helium Gas Pycnometer (Micromeritics AccuPyc II 1340, USA). The true density of the UHMWPE powder was determined to be 0.954 g/cm³. The total, open and closed porosities were calculated using the following equations:=818(2)𝐓𝐓𝐓𝐓𝐓𝐓𝐓𝐓𝐁𝐁 𝐩𝐩𝐓𝐓𝐩𝐩𝐓𝐓𝐩𝐩𝐩𝐩𝐓𝐓𝐩𝐩 = 𝟏𝟏 − (𝛒𝛒𝑩𝑩𝑩𝑩𝑩𝑩𝑩𝑩 𝛒𝛒𝑻𝑻𝑻𝑻𝑩𝑩𝑻𝑻) × 𝟏𝟏𝟏𝟏𝟏𝟏(3)𝐎𝐎𝐩𝐩𝐎𝐎𝐎𝐎 𝐩𝐩𝐓𝐓𝐩𝐩𝐓𝐓𝐩𝐩𝐩𝐩𝐓𝐓𝐩𝐩 = 𝟏𝟏 − (𝛒𝛒𝑩𝑩𝑩𝑩𝑩𝑩𝑩𝑩 𝛒𝛒𝑺𝑺𝑩𝑩𝑻𝑻𝑩𝑩𝑻𝑻𝑺𝑺𝑺𝑺𝑩𝑩) × 𝟏𝟏𝟏𝟏𝟏𝟏(4)𝐂𝐂𝐁𝐁𝐓𝐓𝐩𝐩𝐎𝐎𝐂𝐂 𝐩𝐩𝐓𝐓𝐩𝐩𝐓𝐓𝐩𝐩𝐩𝐩𝐓𝐓𝐩𝐩 = 𝐓𝐓𝐓𝐓𝐓𝐓𝐓𝐓𝐁𝐁 𝐩𝐩𝐓𝐓𝐩𝐩𝐓𝐓𝐩𝐩𝐩𝐩𝐓𝐓𝐩𝐩 − 𝐎𝐎𝐩𝐩𝐎𝐎𝐎𝐎 𝐩𝐩𝐓𝐓𝐩𝐩𝐓𝐓𝐩𝐩𝐩𝐩𝐓𝐓𝐩𝐩3.1.2. Micro-CTThe micro-CT technique is described by Ho and Hutmacher [13]. The internal microstructure of the laser sintered UHMWPE parts was investigated using Skyscan 1172 (Brucker μCT, Kontich, Belgium). The scanning conditions used were a current of 165 µA, voltage of 40 kV, pixel size of 5.9 µm, 360° rotation and 0.35° rotation step.

The scanning conditions used were a current of 165 µA, voltage of 40 kV, pixel size of 5.9 µm, 360° rotation and 0.35° rotation step. SkyScan NRecon software (Bruker micro-CT, Kontich, Belgium) was used to reconstruct cross-section images from tomography projection images obtained from the scanner. CTAn software (Bruker micro-CT, Kontich, Belgium) was used for 3D morphological analysis of the micro-CT data (Figure 4).Figure 4: micro-CT images of LS part at 6 watts: a) Single 2D cross-section image (Pores in black colour), b) Reconstructed images to form 3D part. Red line indicates the position of the 2D image.A cylindrical region of interest (ROI) with a diameter of 5 mm and a height of 8 mm was chosen for the 3D porosity analysis for all samples (Figure 5).

CTVox software was used to show 3D representations of the structure.819Figure 5: 3D representation of the part structure produced by CTVox a) ROI created by CTAn b) Cross-section of the cylindrical ROI (Pores in black colour)3.2. Mechanical propertiesFlexural properties were measured using three-point bending test on Texture Analyser TA500 machine (TM Lloyd Instruments, UK) fitted with a 500 N load cell. The support span length was 40 mm with a constant cross head speed of 2 mm/min. The results of the flexural tests were obtained by averaging the values over three test samples. All tests were performed at ambient temperature.3.3. Scanning electron microscopy (SEM)The external microstructure of the sintered parts was observed using a scanning electron microscope (Philips XL-20, Holland) at an accelerating voltage of 15 kV. All samples were gold sputtered before the observation to avoid charging.4. Results and discussion4.1.

All samples were gold sputtered before the observation to avoid charging.4. Results and discussion4.1. Morphological propertiesMicro-CT technique was used on the UHMWPE porosity samples. Figures 6 and 7 show an image of a whole scanned sample and 2D images of cross-section respectively produced using CTAn software with regions of different densities in different colours. These images show that the density is varied from region to region in the 3D structure. Higher density areas are presented in a white and areas with low density including the pores are presented in black colour.820Figure 6: Reconstructed image of laser sintered sample produced by CTAn softwareAs it can be seen in Figure 6, the results indicate a different density on the inside and in the outer area of the sintered part. The outer shell is more dense (yellow colour) than the inside of the part which is mainly in red to orange colour.

The outer shell is more dense (yellow colour) than the inside of the part which is mainly in red to orange colour. Rüsenberg [14] indicated that the dense shell surrounding the part can be formed and strongly influenced by the thickness of the sample and the cooling rate. Ajuku [15] also reported the same phenomenon and suggested that the short bursts of energy at the edges result in parts being much denser at the edges than at the centres.Figure 7: 2D images of cross section of UHMWPE parts produced at different laser powers821Figure 7 shows the cross-section images of the sintered parts at different laser powers. The dominated colour mainly ranges between red and orange, which indicates a low and medium density. High volume of pores can be observed at all levels of laser power (pores are in black colour).

The dominated colour mainly ranges between red and orange, which indicates a low and medium density. High volume of pores can be observed at all levels of laser power (pores are in black colour). However, there is a fair amount of yellow dots can be seen within the structure as an indication of high density elements.Morphological properties of the 3D structure of the UHMWPE LS parts, such as porosity, average pore diameter and pore size distribution were measured using micro-CT. In this study, two specimens were scanned for each laser power and the average results were obtained.

In this study, two specimens were scanned for each laser power and the average results were obtained. The effect of laser power on porosity of the sintered parts measured by helium gas pycnometer and micro-CT is shown in Table 1.Table 1: Porosity measurement of UHMWPE LS partsHelium PycnometerMicro-CTLaser Power (watts)Total porosity (%)Open pores (%)Closed pores (%)Total porosity (%)Open pores (%)Closed pores (%)664.8564.420.4361.4461.430.01861.4160.980.4360.3060.290.011060.7260.300.4260.6760.660.011262.7962.360.4361.8461.830.01Figures 8, 9 and 10 show the average effect of laser power upon the total, open and closedporosities.822Figure 8: Average effect of laser power upon the porosity of the UHMWPE partsA slight decrease of these porosities can be observed when the laser power increased from 6 to 10 watts. However, further increase in laser power (i.e.

However, further increase in laser power (i.e. 12 watts) showed a slight increase in porosity. Gill et al [16] reported that polyamide LS parts built at higher energy density showed a reduction in porosity when compared to LS parts built at lower energy density. With increasing energy density from the optimum, the porosity increases due to degradation of the polymer.823Figure 9: Average open porosity in UHMWPE parts produced at different laser powersThe change in the trend of total and open porosities measured by micro-CT technique is similar to the porosities measured using helium gas pycnometer method. There are no significant differences in the results obtained from both techniques and it is clear that the results fit very well with each other demonstrating the reliability of micro CT analysis.824Figure 10: Average closed porosity in UHMWPE parts produced at different laser powersThe closed porosity (Figure 10) remains fairly constant with increasing laser power.

However, this result does not fit well with the result of the closed porosity obtained by micro-CT technique. This is very likely due to the fact that the volume of the closed porosity is low. Additionally, if the closed pores are in the same range as the scanning resolution of the micro-CT images, they may not be detected.Figure 11 shows the effect of laser power on the average pore diameter. UHMWPE structure has a slightly larger pores at a laser power of 6 watts and the average pore diameter remains fairly constant with increasing laser power (i.e.

UHMWPE structure has a slightly larger pores at a laser power of 6 watts and the average pore diameter remains fairly constant with increasing laser power (i.e. 8, 10 and 12 watts).825Figure 11: Average pore diameter of UHMWPE parts produced at different laser powersThe cross-section (5 mm diameter) of the 2D micro-CT images in Figure 12 shows a similarcharacteristics at all levels of laser power.Figure 12: Cross-section of micro-CT binarised images of UHMWPE parts produced at different laser powers (Pores in black colour)Figure 13 shows the pore diameter distributions of LS parts at different laser powers. The volume of the pores with diameters in range between 30 – 40 µm is the highest for all levels of826laser power. Pore diameters larger than 40 and smaller than 89 are also present at all laser power levels and pore diameter larger than 89 µm can be seen clearly at laser power of 6 watts.

Pore diameters larger than 40 and smaller than 89 are also present at all laser power levels and pore diameter larger than 89 µm can be seen clearly at laser power of 6 watts. However, the result shows that there is no statistically significant difference in the pore size distribution when the laser power is changed.Figure 13: Pore size distribution of UHMWPE parts produced at different laser powers8274.2. Mechanical propertiesFlexural stress-strain curves obtained by three point bend test analysis are shown in Figure14.Figure 14: Flexural stress-strain curve of UHMWPE parts produced at different laser powersThe result shows a less ductile behaviour for parts built at low laser powers than the parts built at higher laser power. The bonds between the powder particles, which occur due to the formation of more necking, become stronger at higher laser powers, which leads to a more ductile behaviour with a large plastic region in the stress–strain curve.

This behaviour has been observed with Laser Sintering of other polymers [17]. However, at higher laser power, 12 watts, showed a decrease in maximum flexural stress and strain. Gill et al [16] observed this phenomena and reported that with increasing energy density from the optimum, degradation of the polymer occurs, which results in a reduction in strength.Figure 15 shows the result of the flexural stress which was plotted versus the porosity ofthe LS parts.828Figure 15: Flexural stress vs total porosity of UHMWPE LS partsThe porosity results, obtained by the helium gas pycnometer method, show a typical change trend whereby the strength decreases as porosity increases. This behaviour was observed by Gill et al [16] and reported that the strength of LS parts decreases when the laser energy is insufficient to fully sinter the polymer, resulting in an increase in porosity.4.3.

Microstructure observationFigures 16 and 17 show representative SEM images of the surface of the UHMWPEsintered parts at laser power of 6 and 8 watts respectively.829Figure 16: SEM micrographs of the sintered part of UHMWPE at laser power of 6 watts with magnifications: a) 100x, b) 200x, c) 800x and (d) 1600xFigure 17: SEM micrographs of the sintered part of UHMWPE at laser power of 8 watts with magnifications: a) 100x, b) 200x, c) 800x and (d) 1600xThe porous structure can be seen in both samples but with different forms. The SEM observation showed that when the laser power increased from 6 to 8 watts, the UHMWPE LS parts became denser with a higher degree of coalescence. It seems that the fusion of the polymer particles830becomes better with a more compact morphology when the laser power increases to a certain levels [18].5.

It seems that the fusion of the polymer particles830becomes better with a more compact morphology when the laser power increases to a certain levels [18].5. ConclusionsThe following conclusions are made from this study:The results showed a high level of porosity in the UHMWPE LS parts with a range of 60-65% measured by micro-CT and helium gas pycnometer method respectively. A slight decrease in porosity was observed when the laser power increased from 6 to 10 watts. Further increase in laser power to 12 watts showed a slight increase in porosity.The change in the trend of total and open porosities measured by micro-CT technique is similar to the porosities measured using helium gas pycnometer. There are no significant differences in the results obtained from both techniques.

There are no significant differences in the results obtained from both techniques. The results fit very well with each other demonstrating the suitability and reliability of micro CT for morphology characterisation.The UHMWPE samples have slightly larger pores at a laser power of 6 watts and the average pore diameter remains fairly constant with the increase in laser power (i.e. 8, 10 and 12 watts).The flexural stress-strain result revealed a less ductile behaviour for parts built at low laser powers than the parts built at higher laser power. However, at higher laser power (i.e. 12 watts) a decrease in maximum flexural stress and strain was observed.The porosity results, obtained by helium gas pycnometer, combined with the flexural measurements, showed that the flexural strength decreases as the porosity increases.AcknowledgementsThe authors would like to thank The Mellanby Centre for Bone Research at the University of Sheffield and Ms.

Kirsty Nicholson for their technical assistance with the micro-CT. The authors also would like to acknowledge the research funding from the Engineering and Physical Sciences Research Council (EPSRC) of Great Britain (grant number EP/M507611/1). Further contribution from the industrial sponsor (Unilever plc, UK) is also acknowledged with gratitude.831References1.Kerckhofs, G., Pyka, G., Loeckx, D., Van Bael, S., Schrooten, J., and Wevers, M. The combined use of micro-CT imaging, in-situ loading and non-rigid image registration for 3D experimental local strain mapping on porous bone tissue engineering scaffolds under compressive loading.2.De Chiffre, L., Carmignato, S., Kruth, J.P., Schmitt, R., and Weckenmann, A., Industrial applications of computed tomography. CIRP Annals - Manufacturing Technology, 2014. 63(2): p. 655-677.3.Bourell, D.L., Watt, T.J., Leigh, D.K., and Fulcher, B., Performance Limitations in Polymer Laser Sintering. Physics Procedia, 2014. 56(0): p.

fluence of Laser Processing Parameters on the Corrosion Behavior in 316L Stainless Steel Laser Powder DepositionsEswar R. Yarrapareddya, Alan J. Anderson, James W. SearsAdvanced Materials Processing (AMP) Center, South Dakota School of Mines and Technology, Rapid City, SD 57701, USA1. Abstract Laser Powder Deposition (LPD) is an advanced manufacturing process where metallic and/or ceramic powders are injected into a small molten pool produced by a focused laser beam. A computer-controlled robot controls the position of the molten pool, which allows for sequential layering that produces solid free form shapes directly from a CAD/CAM interface. The process parameters are implemented to control the molten pool size by changing the laser processing parameters (i.e. laser power, traverse speed and layer thickness). This work evaluated the corrosion behavior of laser powder deposited 316L stainless steels deposited at different laser processing parameters.

laser power, traverse speed and layer thickness). This work evaluated the corrosion behavior of laser powder deposited 316L stainless steels deposited at different laser processing parameters. The susceptibility of intergranular corrosion (IGC) of laser deposited 316L stainless steel bulk builds is studied by hot boiling sulfuric acid sensitization test. The pitting corrosion behavior of laser deposited 316L stainless steel thin walls is studied by cyclic potentiodynamic polarization methods. From this, the corrosion potential, and current density are calculated. These are related to laser processing parameters. The wrought 316L stainless steel is also tested for comparative study with laser deposited specimens. The laser deposited specimens are about 30% more corrosion resistant than the wrought materials.2.

The laser deposited specimens are about 30% more corrosion resistant than the wrought materials.2. Introduction Laser Powder Deposition (LPD) is an advanced manufacturing technology used to fabricate near-net shape, fully dense metal components directly from a CAD/CAM interface without using traditional machining processes. The LPD system at Advanced Materials Processing (AMP) Center of South Dakota School of Mines and Technology (SDSM&T) consists of a high-power Neodymium Yttrium Aluminum Garnet (Nd: YAG) laser, a Fanuc M-16i six axis computer- controlled robot and two powder feeder units. The powder delivery nozzles are arranged to inject the powder streams directly into the focused laser beam. A schematic representation of LPD system is shown in Figure 1. The system has a flat metal substrate generally made of the metal to be deposited.

A schematic representation of LPD system is shown in Figure 1. The system has a flat metal substrate generally made of the metal to be deposited. The laser beam is focused onto the metal substrate to create a molten pool and the powder delivery nozzles deliver the metallic or ceramic powders into the molten pool to increase the material volume. The computer-controlled robot controls the position of the molten pool. The system repeats the process over and over to deposit a layer of the structure. After one layer is deposited, another is deposited on top of it and so on. The system creates layer upon layer until it has produced a metal version of the CAD model1. The movement of the laser controls the solidification of the molten pool. Typically the solidification rates are high, which produces fully dense materials with excellent mechanical and metallurgical properties.

Typically the solidification rates are high, which produces fully dense materials with excellent mechanical and metallurgical properties. LPD has been to shown to produce fully dense materials with properties equivalent or better than conventionally produced materials2. 116Figure 1. The schematic diagram of laser powder deposition.3. Research Objective The objective of this research is to investigate the corrosion behavior of laser powder deposited 316L stainless steels at different laser-processing parameters. The investigated corrosions are intergranular corrosion and pitting corrosion. These corrosion behaviors are studied in terms of laser-processing parameters. The parameters of interest are laser power, layer thickness, hatch width and linear velocity. For the current work only the laser power is varied, the layer thickness, hatch width and linear velocity are kept constant for all depositions.

For the current work only the laser power is varied, the layer thickness, hatch width and linear velocity are kept constant for all depositions. The laser processing parameters are interrelated, and this relationship is given by the specific energy as defined in the following equation. Specific Energy, SE = (PL x α) / (BR max) Where PL is laser powerα is absorption coefficientBR max is built rateTls is layer thicknessHhs is hatch widthVl is linear velocitySpecific energy, as previously shown by Sears and Marchione, has an effect on the mechanical properties and microstructure of laser powder deposited 690N2 material2. In the current study the corrosion behavior is studied as function of specific energy, but since only the laser power is varied, the results are presented as a function of laser power only. However, it is important to note that this directly relates to specific energy in this case. The computer-controlled robot controls the position of the molten pool.

However, it is important to note that this directly relates to specific energy in this case. The computer-controlled robot controls the position of the molten pool. The control of the molten pool size is predominately117influenced by the absorbed energy from the laser3. The chemical compositions of metallic powders used for depositing 316L SS and wrought 316L SS materials are given in Table 1.Table 1.The chemical compositions of LPD metal powder and 316L wrought material FeCSiMnPSCrNi MoLPD-Metal Powder0.0190.501.440.0220.00716.510.202.09BalWrought316L SS0.031.002.000.0450.00318-2012.002.00Bal4. Laser Powder Deposition Experimental Work The 316L stainless steel (SS) bulk builds and thin walls are laser deposited by using Trumpf 3006L, 3 kW, Continuous Wave, and Nd: YAG laser. A Fanuc M-16i six-axis computer- controlled robot controls the laser-powder head motion.

A Fanuc M-16i six-axis computer- controlled robot controls the laser-powder head motion. Two sample configurations are deposited: a bulk build and a thin wall build (see Figure 2 and 3). Four samples of each are made using the parameters in Tables 2 and 3. The next part describes the corrosion tests used for this study, hot boiling sulfuric acid test is selected for detecting susceptibility to IGC and cyclic potentiodynamic polarization test is used to measure the resistance to pitting. Wrought 316L stainless steel material is also tested for comparative study. The laser powder deposited bulk builds are used for IGC testing and laser powder deposited thin walls are used for generating cyclic potentiodynamic polarization curves.Table 2.

Examples of three builds along with the powder delivery nozzle aligned withthe focused laser beam.Table 3. LPD processing parameters for 316L SS thin wallsParameterSample 1Sample 2Sample 3Sample 4Linear Velocity (mm/s)21212121Layer Thickness (mm)0.500.500.500.50Build Rate (cc/hr)28.428.428.428.4Laser Power (watts)450600750900Figure 3. Example of 900 watts LPD processed thin wall structure.1195. Experimental Procedures5.1 Hot Boiling Sulfuric Acid Sensitization Test Four specimens (550, 650, 750 and 850 watts laser powder deposited) are sectioned from the top portion of laser powder deposited 316L SS bulk builds and a piece of 316L SS wrought material from 316L SS wrought material. All specimens are finished using different grit abrasive papers first, then fine polished with 1 µm and 0.05 µm abrasive solutions. The polishing is to remove any loose oxide layers that might be present on the surface of the specimens.

The polishing is to remove any loose oxide layers that might be present on the surface of the specimens. The total surface area of these specimens is calculated including the edges. The polished specimens are placed in 1000 ml Erlenmeyer flask containing 600 ml of 50% reagent grade sulfuric acid and 25 grams reagent grade ferric sulfate. These specimens are heated for 120 hours as per ASTM G28-97 standard method4. The experimental set up of hot boiling sulfuric acid sensitization test is shown in the Figure 4.Glass CondenserWater Outlet Water InletErlenmeyer flaskSOH 24+Fe2(SO) 34Test specimenHot plateFigure 4.The schematic representation of hot boiling sulfuric acid sensitization test5.2 Cyclic Potentiodynamic Polarization TestLaser powder deposited thin walls (600, 750 and 900 watts) are sectioned into 30x30 mm specimens. A 30x30 mm wrought 316L SS is also sectioned from 316L SS wrought plate. These specimens are first rough polished then fine polished.

A 30x30 mm wrought 316L SS is also sectioned from 316L SS wrought plate. These specimens are first rough polished then fine polished. The polishing is intended to remove any loose oxide layer that might form on the surface and to minimize the effect of surface roughness5. Potentiodynamic polarization tests are conducted on computer-controlled model 273 EG&G potentiostat as per ASTM G 61-86 standard test method6 to study the pitting corrosion behavior. Polarization curves are generated in 3.5% deaerated NaCl solution at room temperature. The NaCl solution is deaerated by purging nitrogen gas for about one hour. The potentials are measured with respect to standard saturated calomel electrode (ESCE=0.242 Volts). Sintered platinum electrode is used as a counter electrode and the test sample served as working electrode. 5 mV/s scan rate is used during the polarization tests. The specimens are immersed in120the test solution for one hour before the polarization.

5 mV/s scan rate is used during the polarization tests. The specimens are immersed in120the test solution for one hour before the polarization. The potential, total current and time are continuously monitored. From these polarization curves the corrosion potential, pitting potential and current density are calculated. The experimental set up of potentiodynamic polarization test is shown in the Figure 5. All experiments are conducted at an ambient temperature of 240C.Figure 5. Potentiodynamic polarization test experimental set up.6.

All experiments are conducted at an ambient temperature of 240C.Figure 5. Potentiodynamic polarization test experimental set up.6. Results and Discussions6.1 Hot Boiling Sulfuric Acid Sensitization Test The effect of the acid solution on specimens was calculated using the following formula6.Corrosion Rate = (K x W) / (A x T x D)Where K = A Constant = 3.45x 106T = Time of exposure, h, to the nearest 0.01 h A = Area, cm2, to the nearest 0.01 cm2 W = Weight loss, g, to the nearest 0.001g and D = Density, g/cm3The measured weight loss and the corrosion rate values are summarized in the Table 4. The presence of IGC is determined by comparing the measured corrosion rate values. The specimens deposited at higher laser power are more IGC resistant than the specimens deposited at lesser laser power.

The specimens deposited at higher laser power are more IGC resistant than the specimens deposited at lesser laser power. 650, 750, 850 watts laser deposited specimens corrosion rate values are lying in the good relative corrosion resistance range and 550 watts laser deposited sample and wrought 316L SS sample lying in the fair range as shown in the Table 5. The relationship between the corrosion rate values and laser power are plotted in the Figure 6. The intergranular corrosion attack can be seen in between the layers in laser deposited specimens and around the grain boundaries in wrought material. The IGC attack in between the grain boundaries in laser deposited specimens and wrought materials is shown in SEM microstructures in Figure 7. The laser powder deposited specimens are about 30% more corrosion resistant than the wrought materials.121Table 4.

Relationship between corrosion rates vs. laser power.122IGCIGCABIGCIGCCDIGCEFigure 7. SEM microstructures of laser deposited 316L SS specimens of laser power 550 (A), 650 (B), 750 (C) and 850 (D) watts and wrought 316L SS (E) after hot boiling sulfuric acid test.123In stainless steels the IGC is mainly due to the impoverishment or depletion of chromium in the grain-boundary areas10. The addition of chromium to ordinary steel imparts corrosion resistance to the steel in many environments. Generally more than 10% chromium is needed to make stainless steel. If the chromium is effectively lowered, the relatively poor corrosion resistance of ordinary steel is reproduced. The carbon diffuses towards the grain boundary quite faster at sensitizing temperatures, but chromium is much lesser mobile. The 17% chromium that is present in 316L stainless steels will be depleted in grain boundary regions resulting in the intergranular corrosion7.

The 17% chromium that is present in 316L stainless steels will be depleted in grain boundary regions resulting in the intergranular corrosion7. In the case of 316L SS, there is very low carbon, so sensitization is generally not a problem; thus the improvement in the LPD specimens is due to another effect. It is hypothesized that the high cooling rates results in a fine microstructure with very little segregation and improved IGC resistance. It is noted that the lower laser powers result in finer microstructure and many more grain boundaries. It is expected that the number of grain boundaries will decrease as laser power increases resulting in improved IGC resistance. It is also hypothesized that the ferrite content decreases as laser power increases thereby decreasing the susceptibility to IGC.6.2 Cyclic Potentiodynamic Polarization TestThe potentiodynamic polarization curves for the laser deposited 316L SS and 316L SS wrought material are shown in the Figure 8.

These curves clearly indicate the specimens deposited at higher laser power have higher corrosion potentials and the specimens deposited at lesser power have lesser corrosion potentials. The corrosion potentials for different specimens are summarized in the Table 6. The current densities are calculated by drawing tangents to the anodic and cathodic curves, where there is a sharp increase in the potential, and extrapolating the tangent’s intersection point onto the x-axis. The calculated current densities are summarized in Table 7. These show the lower laser power deposited specimens are less pitting corrosion resistant and higher laser power deposited specimens are more pitting corrosion resistant.The graphs showing the relationship among corrosion potential, current density and laser power are shown in the Figures 8 and 9. The potential at which the current sharply increases is defined as the pitting potential.

The potential at which the current sharply increases is defined as the pitting potential. The potential up to which we can protect the material from pitting is called protection potential. If the pitting potential and protection potential are the same, there will be little tendency to pit. If the protection potential is more positive (anodic) than the pitting potential, there will be no tendency to pit. If the protection potential is more negative than the pitting potential, pitting could occur8. There might be two possible reasons for the improved pitting resistance after laser powder deposition: (i) less segregation the major alloying elements Cr and Ni, or (ii) suppression or elimination of inclusions by high cooling rates9.In summary, potentiodynamic polarization curves are useful in studying the localized corrosion such as pitting corrosion. This describes how material will behave when exposed to particular liquid environment.

The relationship between corrosion potential and laser power.125Table 7. Current densities for different laser deposited specimensLaser Power, wattsCurrent Density, nA/cm2Wrought 316L SS60360050175035590089Current Density vs Laser Power700600Wrought) 2 m c / a n ( y t i s n e D500400300750600t n e r r u C2001009000012345S ample NumberFigure 10. The relationship between current density and laser power.7. Conclusions and Future work7.1 Conclusions and DiscussionsThe laser powder deposited 316L stainless steels with higher laser power are more IGC resistant than the specimens deposited at lesser laser power and 316L stainless steel wrought materials. In fact, the 850 watts laser deposited specimen is about 30% more IGC resistant than 316L stainless steel wrought materials.

In fact, the 850 watts laser deposited specimen is about 30% more IGC resistant than 316L stainless steel wrought materials. The possible reasons for the improved IGC resistance might be that as the laser power increases grain size also increases and lesser total grain boundary area will be available for IGC attack. Another possible reason might be, as the laser power increases there might be decrease in the percentage of ferrite content in final product, if we assume the ferrite phase plays major role in IGC attack. It is also hypothesized that the high cooling rates results in a fine microstructure with very little segregation and improved IGC resistance. It is noted that the lower laser powers result in finer microstructure and many more grain boundaries. It is also126expected that the number of grain boundaries will decrease as laser power increases resulting in improved IGC resistance.

It is also126expected that the number of grain boundaries will decrease as laser power increases resulting in improved IGC resistance. The corrosion potentials of laser powder deposited specimens are moving more towards the noble direction as the laser power is increasing. The pitting corrosion resistance of laser powder deposited specimens deposited at higher laser power is higher than that of laser powder deposited specimens deposited at lesser laser power and wrought material. In summary, the laser powder deposited specimens give about 30% improvement in corrosion rate over the conventional wrought materials. This would demonstrate that the laser powder deposition gives improved corrosion resistance materials over conventional wrought materials. This improvement in the corrosion resistance will improve the longevity of the materials used for different applications.

This improvement in the corrosion resistance will improve the longevity of the materials used for different applications. The current results are in agreement with Optomec LENS-deposited 420 stainless steel results10.7.2 Future WorkThe future work may include:1. X-ray diffraction analysis to identify and quantify the phases present before and after testing specimens for hot boiling and potentiodynamic polarization tests. This work will tell us what phase is mainly corroding.2. Quantify the grain boundary surface area to show that it changes with laser power. This will support the hypothesis that some of the increase in corrosion resistance is caused by the decrease in grain boundary surface area.3. Investigating the other types of corrosion behavior of laser deposited specimens and comparing the results with wrought materials.8. AcknowledgementsThis work is fully supported by U.S. Army Research Laboratory and U.S. Army Office under the grant number DAAD 19-02-2-001.

AcknowledgementsThis work is fully supported by U.S. Army Research Laboratory and U.S. Army Office under the grant number DAAD 19-02-2-001. Support from the infrastructure of AMP Center and Department of Materials and Metallurgical Engineering of SDSM&T is also acknowledged. The authors would also like to thank Mr. Bill Arbegast and Dr. Glen Stone for their invaluable suggestions.9. References1. Richard Grylls., “Laser Engineered Net Shapes” in Advanced Materials & Processes/ January 2003, Volume 161, No.1, PP 45-46.2. Sears, James W. and Marchione, Thierry, “The Effects of Processing Parameters on Microstructure and Properties of Laser Deposited PM Alloy 690N2 Powder”, 2002 International Conference on Metal Powder Deposition for Rapid Manufacturing.3. Sears, James W.

The vertically printed interface was found to have a smooth surface, while the horizontally printed interface had a micro-zigzag interlocking geometry. The formation of this micro-zigzag interlock is due to the switching of extruders during printing, which mechanically reinforces the interface adhesion strength. Using butt-joint tensile test and microscope observation, it was found that the geometry of this zigzag interlock is significantly influenced by the layer thickness, nozzle temperature and extruder offset.

By optimizing the layer thickness and nozzle temperature, the interface adhesion strength between dissimilar materials was increased by 58.2% without significantly increasing the printing time or fabrication complexity.Keywords: Fused deposition modeling (FDM), Interface adhesion strength, Microstructure, ProcessingIntroductionThe development of multi-material additive manufacturing(AM) technologies allows the fabrication of multi-functional structures by printing different materials into the same component [1]. Nowadays, one of the most widespread MMAM technologies is multi-material fused deposition modeling (MMFDM), which is usually implemented by equipping FDM machines with multiple extrusion nozzles [2].

Nowadays, one of the most widespread MMAM technologies is multi-material fused deposition modeling (MMFDM), which is usually implemented by equipping FDM machines with multiple extrusion nozzles [2]. However, achieving sufficient interfacial bonding strength between materials with dissimilar chemical properties is always a challenge in MMFDM due to its discontinuous nature: a molten filament is extruded and deposited onto the previously deposited layer, forming bonds with adjacent filaments.In mono-material parts, the formation of bonding between adjacent filaments was divided into three processes based on the polymer welding theory: 1) surface contacting; 2) neck growth; 3) molecular diffusion at interface and randomization (entanglement) across the inter-filament interface [3]–[5].

During printing, coalescence of neighbor filaments starts instantaneously after surface contact, resulting in the growth of neck length, where surface tension is the driven force and viscous and inertial forces are resistant forces [4], [6]. Meanwhile, the interfacial bond strength continues to grow as polymer chains diffuse across the contacted interface until T < Tg [6]. Many1030studies concluded the weakness of mono-material interfacial bond strength in FDM to the insufficient interfacial molecule diffusion [6]. While other studies arguing that microstructure of the interlayer contact rather than incomplete molecular bonding is the predominant cause of strength anisotropy in FDM [7]. It’s reported that for a range of materials, printing conditions and microscale geometries, bulk-material strength can be achieved by taking actual geometry of interface into strength calculation[8].

It’s reported that for a range of materials, printing conditions and microscale geometries, bulk-material strength can be achieved by taking actual geometry of interface into strength calculation[8]. A comprehensive model which combined interlayer contact model with a diffusion model was developed by Coogan and Kazmer [9] to predict the interlayer strength. The result suggested that the interfacial bond strength of acrylonitrile butadiene styrene (ABS) parts was governed by limited interlayer diffusion due to the long relaxation time of ABS, while the high-impact polystyrene with shorter relaxation time had complete diffusion and was constrained by a lack of interlayer contact.The formation of multi-material interfacial bonding can also be described by the polymer weld model but is influenced by many other factors. Firstly, the materials at interface can have very different chemical properties.

Firstly, the materials at interface can have very different chemical properties. The mechanical integrity of the interface is expected to be reduced when the different printed materials have low chemical affinity, such as polylactic Acid(PLA) and thermoplastic polyurethane(TPU) [10]. Secondly, the switch of extruders leads to additional time of cooling which changes the thermal history of printing. Finally, the multi- extruder deposition process introduces extra geometrical errors during printing [10]. No matter how carefully the alignment process is done, there will always be offsets between these extruders, leading to the separation or overlap of printed components. New experiments and models should be developed to take the effects of these new factors into account.Recently, the effect of process parameters on the multi-material interfacial strength was studied in several research. Yin et al.

Yin et al. investigated the effect of nozzle temperature, bed temperature and print speed on the adhesion strength of ABS-TPU interface. It’s found that the adhesion strength can be doubled by increasing the bed temperature from 30 to 68°C. Tamburrino et al. [11] reported that the decrease of infill ratio has a negative effect on the multi-material interface adhesion strength. Other studies devoted to affix materials with weak chemical bond with 2D or 3D mechanical interlocking such as T-bones and dovetail interlocks [12], interlaced topologically interlocking lattice [13] and triply periodic minimal surface [14].

These macro-scale mechanical interlocks significantly increase the adhesion strength between dissimilar materials but make the manufacturing process more complex.Although the effect of microstructure on the adhesion strength of mono-material interfaces has been shown to be significant, its effect on the strength of a multi-material interface has not been studied in the literature. This study aims to investigate the effect of microstructure on the interface adhesion strength between materials with dissimilar physical and chemical properties. The effect of two key process parameters, namely nozzle temperature and layer thickness, was investigated.

The effect of two key process parameters, namely nozzle temperature and layer thickness, was investigated. Butt-joint tensile test and microscopic observation were performed to evaluate the adhesion strength of the printed multi-material interfaces.Materials and methodsIn this study, the interface adhesion strength between a typical rigid material, PLA [15] and a semi-flexible material, TPU [16], were investigated using butt joint tensile test and1031microscope. The combination of these two materials offers a possibility to develop functional structures. However, their vastly different physical and chemical properties make achieving sufficient interfacial adhesion strength a challenge.Experiment and specimen designA universal tensile testing machine was used to study the tensile strength of the PLA-TPU interface. All tests were performed at a loading speed of 5 mm/min.

All tests were performed at a loading speed of 5 mm/min. The test length between the wedge screw grips was 40 mm and the measured length between the extensometers was 25mm (Figure. 1(a)). To avoid the influence of extruder offset on the interface strength, a horizontally printed dual-interface specimen was designed so that the specimen always fractured at the weaker interface (Figure. 1(b)). Red stands for PLA and grey stands for TPU. A vertically printed single- interface butt-joint specimen was used as reference specimen. The total length, width and thickness of these two specimens were the same. As demonstrated in Figure. 1(c) and (d), the geometry of printed interface is controlled by layer thickness (h) and extrusion width (w). The interfacial geometry is supposed to be significantly influenced by the build orientation even if layer thickness and extrusion width are the same.Figure. 1 Fabrication and butt-joint tensile test of specimens.

The layer thickness (h) controls the geometry of the interface. To investigate the effect of these two process parameters on the tensile strength of PLA-TPU interface, dual-interface specimens were printed at three levels of T and h with 4 repetitions (Table. 1). T stands for the nozzle temperature at which TPU was printed. PLA was always printed at a nozzle temperature of 210°C. 4 single-interface specimens were printed vertically as a reference for the horizontally printed dual-interface specimens. Other process parameters were the same for all the specimens: infill ratio, 100%, raster angle, 90°; extrusion width, 0.4 mm; bed temperature, 60°C; number of contours, two; fan speed, 50%. Each type of specimen was printed with an additional specimen for microscopic observation.Results and discussionsInterface geometry of specimensSpecimens were cut and polished to be observed under optical microscope.

It was found that the horizontally printed PLA – TPU interface had a micro-zigzag interlocking geometry (Figure. 2(a)(b)), while the vertically printed interface had a smooth surface (Figure. 2(c)). Comparing the left and right interface (Figure. 2(a) and (b)) of horizontally printed specimen, it’s found that both of them have zig-zag interlock geometry, while the left interface has a much deeper interlock. This indicates that the actual offset of the TPU extruder was shifted too far to the left. Furthermore, similar micro-zigzag interlocking was also found in the horizontally printed PLA- PLA interface (Figure. 2(d)). This suggests that this zigzag interlocking is a general effect of the printing process, independent of the material.Figure. 2 Geometries of printed interfaces.

2(d)). This suggests that this zigzag interlocking is a general effect of the printing process, independent of the material.Figure. 2 Geometries of printed interfaces. (a) Right interface and (b) left interface of horizontally printed dual- interface specimen; (c) interface of vertically printed specimen; (d) interface of PLA-PLA interface; (e) formation process of the zig-zag interlock; (f) Geometry of the PLA-TPU interface.As shown in figure. 2(e), the formation of this interlocking was supposed to be a result of extruder switch and material wetting. To save the printing time, two layers of material are deposited after each extruder switch in a standard multi-material FDM process. Then, cooling fans are used to accelerate the solidification of the deposited beads.

Then, cooling fans are used to accelerate the solidification of the deposited beads. The expansion of the beads deposited in the first layer is constrained by the beads previously deposited by the other extruder.1033However, in the second layer, the bead deposited next to the interface is expanded due to material wetting, during which surface tension is the driven force and viscous and inertial forces are resistant forces. The repetition of this process resulted in the zig-zag interlock interface observed in Figure. 2. In vertically printed specimens, only one time of extruder switch was happed which resulted in a smooth interface.Mechanical response under tensile loading and microstructure of the interfacesFigure. 3 Load-displacement curves, corresponding microstructures and fracture surfaces at interfaces.

3 Load-displacement curves, corresponding microstructures and fracture surfaces at interfaces. (a) Specimens printed at T = 230°C with (1)(4) h = 0.1 mm, (2)(5) h = 0.2 mm, (3)(6) h = 0.3 mm; (b) Specimens printed at T = 240°C with (1)(4) h = 0.1 mm, (2)(5) h = 0.2 mm, (3)(6) h = 0.3 mm; (c) Specimens printed at T = 250°C with (1)(4) h = 0.1 mm, (2)(5) h = 0.2 mm, (3)(6) h = 0.3 mm. Red part is PLA and while part is TPU. The direction of the deposited filaments is orthogonal to the surface of microscopic images.1034Figure. 3 shows the typical load-displacement curves of specimens under quasi-static tensile loading with T = 230°C, 240°C and 250°C. The corresponding microstructures and fracture surfaces of interface were also presented. In general, the interfacial adhesion between PLA and TPU formed upon printing is not strong enough so that all the specimens failed at the interface region. All the specimens have a micro-scale zig-zag interlock geometry at the interface.Figure.

All the specimens have a micro-scale zig-zag interlock geometry at the interface.Figure. 4 shows the interface strength of specimens influenced by T and h. The interface strength was calculated by dividing the maximum tensile load by the cross-sectional area of the specimen (Table. 2). In the cases of T = 230°C and 240°C, the decrease in h from 0.3 to 0.1 mm significantly increases the interface strength (Figure. 4). Particularly, at T = 230°C, the interface strength of specimens printed at h = 0.1 mm is 12 times of that of the specimen printed at h = 0.3 mm. Meanwhile, under microscope, the geometry of interfaces printed at smaller layer thickness have deeper zig-zag geometry and smaller pores. Therefore, the contact length between PLA and TPU is much longer at smaller layer thickness, which results in a stronger mechanical interlock and explains the increase in interface strength.

Therefore, the contact length between PLA and TPU is much longer at smaller layer thickness, which results in a stronger mechanical interlock and explains the increase in interface strength. From another aspect, more residual materials can be found in the fracture surfaces of specimens printed at h = 0.1 mm (Figure. 3(a)(1), (b)(1) and (c)(1)). This indicates that specimens printed at smaller h has stronger zig-zag interlock. However, in the cases of T = 250°C, the interface strength is quasi-constant as h decreases from 0.3 to 0.1 m (Figure. 4). It’s supposed that the viscosity of material is small enough at high nozzle temperature to achieve a complete neck growth at the interface. So, the influence of interlocking is less important.Figure. 4 Tensile strength of specimens.Table.

So, the influence of interlocking is less important.Figure. 4 Tensile strength of specimens.Table. 2 Tensile strength of dual-interface specimensT (°C)230 240 2500.1 4.77 ± 0.10 5.06 ± 0.08 3.69 ± 0.23h (mm)0.2 1.55 ± 0.19 2.68 ± 0.43 2.9 ± 0.690.3 0.39 ± 0.07 2.37 ± 0.61 3.82 ± 0.17In the cases of h = 0.2 mm and 0.3 mm, the increase in T from 230°C to 250°C significantly increases the interface strength (Figure. 4). That’s because the increase in T reduces the viscosity of TPU and improves the neck growth at the interface. As shown in Figure.10353, specimens printed at higher T have smaller pores in microstructure and more residual materials on the fracture surfaces. In the case of h = 0.1 mm, the interface strength slightly increased when T is increased from 230°C to 240°C, but decreases significantly as T is increased from 240°C to 250°C.

In the case of h = 0.1 mm, the interface strength slightly increased when T is increased from 230°C to 240°C, but decreases significantly as T is increased from 240°C to 250°C. This is supposed to be a result of material degradation, as 250°C is too high for the extrusion of TPU (melting temperature of TPU is 180°C).Optimization of manufacturing using varied layer thicknessExcept for several cases, the decrease of h and the increase of T significantly improve the interface adhesion strength between PLA and TPU. However, the decrease of h not only increases the interface adhesion strength but also significantly increases the printing time. To solve this problem, a specimen with varied layer thickness was developed (Figure. 5). The aim is to reinforce the zig-zag interlock effect by using a small h near the PLA-TPU interface, while using a bigger h at other regions to save the printing time.Figure. 5 The optimized specimen with varied layer thickness.

5 The optimized specimen with varied layer thickness. l1 =32mm, l2 = 6 mm; h1 = 0.2 mm, h1 = 0.1 mm.To verify the effectiveness of this method, 4 optimized specimens were printed and tested. As shown in Figure. 6, all the specimens were printed at T =240°C. The first three type of specimens were horizontally printed with different layer thickness. “Varied h” refers to the optimized specimen with varied layer thicknesses. “Ref” is the vertically printed reference specimen with h = 0.2 mm. The results indicate that the interface adhesion strength of the optimized specimen was 58.2% higher (4.24 Mpa) than that of the specimen with a constant h = 0.2 mm (2.68 Mpa). Meanwhile, the printing time was only increased by 28.2% (44.6 min)Figure.

Meanwhile, the printing time was only increased by 28.2% (44.6 min)Figure. 6 Interface adhesion strengths and printing times of the specimens with different layer thickness.Conclusion1036In this study, the interface adhesion strength between a typical rigid material, PLA and a semi-flexible material, TPU, were investigated using butt joint tensile test and microscope. The experiment results reveal a new adhesion mechanism in multi-material FDM: micro-scale zig-zag interlock. It’s found that the vertically printed specimens have a smooth interface, while the horizontally printed specimens have a micro-scale zig-zag interlock interface. The formation of this micro-zigzag interlock is due to the switch of extruders and material wetting.

The formation of this micro-zigzag interlock is due to the switch of extruders and material wetting. The formation of this microscale zig-zag interlock mechanically reinforces the interface adhesion strength, which is firstly reported in the literature.The effect of two key process parameters, namely nozzle temperature and layer thickness, was investigated. In general, the decrease in layer thickness increases the depth of zig-zag interlock and results in longer contact length between PLA and TPU and higher interface adhesion strength. The increase in nozzle temperature improves the material wetting and neck growth at the interface, leading to a stronger interlock and higher interface adhesion strength.

Freeform parts produced by SLS typically remain buried ina bed of loose powder until they have cooled enough to be handled. Non-isothermalconditions result in heat transfer from the part to the surrounding powder which can causesecondary, uncontrolled sintering, or part growth. Part growth (densification) isaccompanied by the formation of gas-filled voids as required by conservation of massprinciples. The extent of growth is dependent upon local sintering rates which increaseexponentially with temperature (Nelson, 1993). Hence, the ability to predict thetemperature distribution induced by various modes of heat transfer might aid in thedevelopment of thermal control techniques to produce accurate (minimal part growth) finalfreeform part shapes.In this paper, a first attempt is made to quantify the influence of surface radiationand natural convection heat transfer on the temperature distribution within a polycarbonatepowder under non-isothermal conditions.

At steady-state, local temperatures and sinteringrates are obtained for cases involving pure conduction within the powder along withradiative and convective cooling of the powder surface. These results are compared to thebase case involving pure conduction heat transfer through the powder and surrounding air.MODELThe 2-dimensional, non-isothermal arrangement of Figure 1 depicts the idealizedscenario under consideration. Two concentric, square, isothermal tubes serve as heatsource and sink. The outer tube is maintained at 180°C while the inner tube is held at 50°C. The powder has an initial porosity of 0.5. The region above the powder surface isoccupied by air. The 2-dimensional, steady-state temperature distribution within both theair and powder is determined using the commercial code FillAP with the followingdescriptive equations.6201..........----25.4 mm - - -......l1l-I1AirT16.93 mm ~ 9 4-8.47 mmlFigure 1.

System configuration.The governing equations for the pure fluid and the saturated porous medium aregiven separately. In writing the following equations it is assumed that fluid flow is laminar,steady, incompressible, and two-dimensional. For the pure fluid region, the twodimensional, steady-state Navier-Stokes equations are used. Conservation of mass is givenbyav au -+-=0 ay axConservation of momentum is described byandav) a ( p u-+v- = - -+ - Jlf- + - Jlf- +pg~(T-Tc) ay ax( au axau) ayap aya ( axav) ayConservation of energy is expressed aspc f(U aT + v aT) = ~(kf aT) + ~(kf aT) ay axaxayaxayP621(1)(2)(3)(4)All thermophysical properties of the pure fluid are assumed constant with the soleexception being density, where a Boussinesq approximation is used.For the saturated porous medium, a Brinkman-Forchheimer extended version of theDarcy model is used, as discussed in Beckermann et al. (1987) and Kaviany (1991).

(1987) and Kaviany (1991). It isimportant to note that the velocities within the saturated powder are Darcian, orvolumetrically-averaged, velocities. Conservation of mass within the powder is given byConservation of momentum within the powder bed, given byandincorporates the Brinkman extension (the second and third terms in the above equations)and the Forchheimer extension (the last term) to account for viscous and inertial effects.Conservation of energy is described byAll thermophysical properties of the saturated powder are assumed to be constant, with theexception being a Boussinesq approximation for the air density. Permeability of the powder was measured to be 1.5xl0-10 m2, as described in Norrell (1996). For the saturated powder fleff and keff are effective air viscosity and thermal conductivity, respectively.

Permeability of the powder was measured to be 1.5xl0-10 m2, as described in Norrell (1996). For the saturated powder fleff and keff are effective air viscosity and thermal conductivity, respectively. Ithas been found previously by Neale and Nader (1974) that using Jleff=Jlf within theBrinkman extension provides good agreement with experimental data. A simple weightedaverage given by622(5)(7)(8)(9)is used for the effective thermal conductivity. The C term in the Forchheimer extension isthe inertia coefficient given in Ergun (1952) asc _ 1.753/2-)175 $Radiation view factors between the (radiatively black) powder and surrounding(black) surfaces are determined using the Hottel crossed-string method to obtain viewfactors at each powder surface node (Seigel and Howell, 1981).

The radiative fluxes onthe exposed powder surface are calculated and mapped into FIDAP as heat flux terms.Interparticle (within powder bed) radiation exchange is neglected here.The model for viscous sintering is obtained from Nelson et al. (1993)d$ / dt = K' .($ - $00 )where the coefficient isThe activation energy (Ea) and other powder properties are also obtained from Nelson et al. (1993). The sintering rate coefficient (A) used in this study is obtained from Kandis andBergman (1996).Solid shapes are estimated based upon the predicted steady state temperaturedistribution in the powder using an original consolidation model which is described indetail in Kandis and Bergman (1996).

In short, the powder domain shape evolves in astress-free manner, with local sintering rates inducing 2-dimensional solid phase velocitieswhich are then integrated in time to yield the final powder region shape.RESULTSThe pure conduction temperature distribution is presented in Figure 2a. Isotherms are separated by 1aoc in all figures. The isotherms in the powder and surrounding air are roughly circular in shape with a discontinuity in slope at the powder surface. The highestsintering rates occur in the hottest regions (near the outer tube). Less sintering takes placenear the central region of the powder surface where temperatures are lowest. The finalpredicted part shape as a result of part growth is shown in Figure 2b. This figure is623(10)(I1a)(lIb)obtained by integrating local sintering rates in time for a period of 6 hours. The partcontains a vertical crack in the center, surrounded by a shallow trough or semi-circularditch.

The sintered part has a solid and smoothtexture along the bottom and sides (regions exposed to the hot outer tube) while the innersurface displays a soft, flaky appearance. Although not discernible from the photograph, afine crack in the center of the trough runs axially along the length of the test specimen.Qualitatively, the part shape compares well with the shape in Figure 4b.Figure 5. Experimentally-grown part shape.CONCLUSIONSThe results presented in this study indicate that, for the material, geometry andthermal conditions considered here, surface radiation heat transfer has little effect on thetemperature distribution and powder sintering. Only a marginal decrease in part growth isobtained with the inclusion of radiation.

Only a marginal decrease in part growth isobtained with the inclusion of radiation. In contrast, natural convection (which may occurin large, buried voids which form in response to bed consolidation associated with partgrowth) can modify temperatures and sintering rates, in tum, affect part growth.Comparison with experimentally grown parts shows the influence of natural convection tobe the more significant factor in determining part shape and growth.REFERENCESBeckermann, C., Ramadhyani, S., and Viskanta, R., 1987, "Natural ConvectionFlow and Heat Transfer Between a Fluid Layer and a Porous Layer Inside a RectangularEnclosure," J. Heat Trans., Vol. 109, pp. 363-370.Ergun, S., 1952, "Fluid Flow Through Packed Columns," Chern. Eng. Prog.,Vol. 48, pp. 89-94.Kandis, M. and Bergman, T. L., 1996, "Void Formation and Crack Propagation inPolymer Parts Grown by Non-Isothermal Sintering," in Proc. ofthe National Heat Trans.Conf, edited by V. Prasad et aI., HTD-Vol.

and Bergman, T. L., 1996, "Void Formation and Crack Propagation inPolymer Parts Grown by Non-Isothermal Sintering," in Proc. ofthe National Heat Trans.Conf, edited by V. Prasad et aI., HTD-Vol. 323, ASME, New York, pp. 199-206.626Kaviany, M., 1991, Principles ofHeat Transfer in Porous Media, Springer-Verlag,New York.Neale, G. and Nader, W., 1974, "Practical Significance of Brinkman's Extensionof Darcy's Law: Coupled Parallel Flows Within a Channel and a Bounding PorousMedium," Canadian J. Chern. Eng., Vol. 52, pp. 475-478.Nelson, J. C., Xue, S., Barlow, J. W., Beaman, J. J., Marcus, H. L. andBourell, D. L., 1993, "Model of Selective Laser Sintering of Bisphenol-A Polycarbonate,"Ind. Eng. Chern. Res., Vol. 32, pp. 2305-2317.Norrell, J., Wood, K., Crawford, R., and Bergman, T., "Forced Convection in aPolymeric Powder," 1996, to be published in Proc. 1996 SFF Symposium, Austin, TX.Seigel, R. and Howell, J.

2305-2317.Norrell, J., Wood, K., Crawford, R., and Bergman, T., "Forced Convection in aPolymeric Powder," 1996, to be published in Proc. 1996 SFF Symposium, Austin, TX.Seigel, R. and Howell, J. R., 1981, Thermal Radiation Heat Transfer, HemispherePublishing Company, U. S. A.627628.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceINFLUENCE OF PART MICROSTRUCTURE ON MECHANICAL PROPERTIES OF PA6X LASER SINTERED SPECIMENSC.(cid:3)Kummert*, W. Diekmann†, K. Tews* and H.J. Schmid *Direct Manufacturing Research Center/ Particle Technology Group, Paderborn University, Paderborn, Germany †Evonik Resource Efficiency GmbH, Marl, GermanyAbstractThe influence of different process parameters on mechanical properties of selectively laser sintered (SLS) parts was investigated in various studies. Until now, the significant differences in mechanical characteristics depending on processing conditions are insufficiently explained but reasons may be found in part microstructure. For this reason, PA6x test specimens with different component properties were printed by changing laser exposure strategies and their microstructure was examined using for example XCT-analysis.

For this reason, PA6x test specimens with different component properties were printed by changing laser exposure strategies and their microstructure was examined using for example XCT-analysis. PA6x is a comparatively new SLS material which offers outstanding mechanical properties if adequate SLS processing parameters are used. In this study different SLS machines are used by EVONIK and the DMRC, Paderborn University to investigate the relationship between SLS-specific manufacturing conditions, the resulting components microstructure and finally the component properties. The knowledge gained can contribute to a deeper understanding of the process.IntroductionSelective laser sintering is a promising additive manufacturing process for serial production because of its high productivity due to the 3-dimensional set up of the components inside the build space. Besides efficiency, high quality and reproducible build parts are required in the industry.

Besides efficiency, high quality and reproducible build parts are required in the industry. Especially new SLS materials, developed to address the demands of the industry covered with PA12, bring new challenges and have not yet been investigated extensively. When a new material for the SLS process is launched on the market, suitable process parameters must be developed first to generate the best possible characteristic values. Not only optimization of build part properties but also improvement of robustness and homogeneity independent of build orientation, positioning or other process specific effects take place in parameter development. The most controllable parameter in the process is the laser exposure pattern.

The most controllable parameter in the process is the laser exposure pattern. The influence of the different variables of the laser energy input like laser power and scan speed on resulting part characteristics has been investigated in different works [5, 7].However, mechanical properties are the result of the part microstructure like porosity, crystallinity, state of the polymer chains and the surface structure. These are also influenced by the original powder properties, such as particle size distribution and crystallization temperature and melt degree [4]. Furthermore, the rheological behavior of the polymers is directly related to the degree of particle melt [2, 6]. A higher melt viscosity requires more laser energy input to melt and coalescence the polymer particles. On the other side, insufficient laser energy leads to higher porosity and therefore reduced mechanical properties [7].Wegner et al. [8] placed greater attention on the latter.

On the other side, insufficient laser energy leads to higher porosity and therefore reduced mechanical properties [7].Wegner et al. [8] placed greater attention on the latter. They showed that a double laser exposure strategy can help to improve elongation at break and to reduce fluctuations due to cycle time as well as build direction and positioning. As long as the energy input is not too high a second exposure reduces pores and thus improves layer bonding.A double exposure of the contour region was already found to generate lower porosity within the part edges and therefore to improve mechanical properties especially in z-build direction [3]. Consequently, it would be interesting to know, if there is a maximum porosity inside of the sample which is not compromising part quality provided by the contour has a dense structure.

Consequently, it would be interesting to know, if there is a maximum porosity inside of the sample which is not compromising part quality provided by the contour has a dense structure. Furthermore, during parameter development for PA613 it could be noticed that despite identical process conditions and parameters elongation at break shows high fluctuations which was already recognized for other materials [6, 8].735When comparing data of injection molded and laser sintered nylon 613 (polyamide 613, PA613) a discrepancy can be noticed (see Table 1). As already known from PA12 PA613 shows also a shift to a slightly higher young’s modulus from about 1900 MPa (molded) to 2000 MPa (printed). Also known from PA12 is that the tensile strength of injection molded and laser sintered parts is comparable. This is also the case for PA613 which shows a tensile strength of about 55 MPa. The mentioned discrepancy can be found in elongation at break (EaB) and notched impact strength (NIS).

This is also the case for PA613 which shows a tensile strength of about 55 MPa. The mentioned discrepancy can be found in elongation at break (EaB) and notched impact strength (NIS). While injection molded PA12 has a higher EaB and NIS than injection molded PA613, the opposite can be observed after laser sintering both nylons, i.e. PA613 now shows a higher EaB and NIS than PA12.Table 1: Data for injection molded and laser sintered parts of PA12 and PA613"' "' iii ., ·;: ., .!! ,a " :. ct0Test methodPropertyYoungs modulusTensile TestingTensile strengthElongation at BreakCharpy testNotched impact strength at RTDensity ArchimedesDensityMicroscopycrystallinityUnitMPaMPa%KJ/m'g/cm'... f 32 0 E C: >< ·u 0 CD f a, s ...

f 32 0 E C: >< ·u 0 CD f a, s ... f C:NC) C:NC) ·c: 2 C: ·;;; <ii >< CD Vl f <ll ...J150019001650200045554555>300>2001540357351,011,0550,95-1 ,001,00-1 ,05Data were generated by building parts with special parameters. The mechanical properties can e.g. vary depending on the part position (SLS), type of test bar and the operating conditions which are used. Injection molded specimen are made with standard settings. Elaborated settings can lead to different results.In order to better understand these effects, the microstructure shall be examined more closely. Porosity, crystallinity and crystalline structure are known to have a high influence on mechanical properties [1, 4]. With increasing crystallinity and ordered lamella in spherulites part strength is pronounced. On the other hand side, elasticity requires amorphous regions [1].

With increasing crystallinity and ordered lamella in spherulites part strength is pronounced. On the other hand side, elasticity requires amorphous regions [1]. To take these microstructural properties into account crystallinity is measured by DSC (differential scanning calorimetry) and spherulites are detected with light microscopy.ContourFillingFigure 1: Cross section of tensile specimenIn the following, laser exposure parameters of the filling and contour (see Figure 1) are varied on two different SLS machine types to build PA613 specimens. Microstructural properties, especially the porosity are correlated with mechanical properties to find the origin of the sensitivity towards process parameters. In this way, the goal of manufacturing parts with high quality and high reproducibility may be achieved.736737medium densities, part density is calculated. help of an analytical balance the weight is determined in air and ethanol.

help of an analytical balance the weight is determined in air and ethanol. Under consideration of the particular dimensions are needed. Therefore, samples are taken of the measuring length and immersed in ethanol. With the the build part density the immersed body method according to DIN EN ISO 1183 is applied as no uniform which is analyzed in detail, was separated into the contour area and filling area as described in [3]. To determine defect sizes are taken into account. Hereby the voxel size is 10 (cid:586)m. Furthermore, the Region-Of-Interest (ROI), from Volume Graphics GmbH. The algorithm compares the gray scale value with a threshold, so that only defined Furthermore, the 3D XCT-data are analyzed using the “Defect Analysis” of the software VGstudio MAX 2.2 X-ray tomography) was performed with GE Phoenix Nanotom S with 70 kV 150 (cid:586)A and 1440 projections. For this purpose a Netzsch DSC 214 Polyma is used with a heating rate of 10 K/min.

For this purpose a Netzsch DSC 214 Polyma is used with a heating rate of 10 K/min. The XCT-scan (computed the crystallinity in the following and is determined by DSC (differential scanning calorimetry) measurements. EH universal testing machine. The fifth specimen is taken for microstructural tests. Melt enthalpy corresponds to At least four specimens for each parameter set are taken for tensile tests, performed on an INSTRON 55692x first0.13000402x first0.353000402x afterwards0.225450040t1x first 1x afterwards0.2256000401x afterwards0.35600025Sequence[mm] Beam offset[mm/s] Scan speed[W] Laser powerTable 2: Contour exposure parameters of the 1st sample series with VESTOSINT® Z2837corresponding design of experiment is shown in Table 2 and Table 3. to reduce the polymer particle coalescence or by a very intense exposure resulting in material evaporation. The dense contour region.

to reduce the polymer particle coalescence or by a very intense exposure resulting in material evaporation. The dense contour region. Deviating from F0 pores are generated by a higher hatch distance and a lower energy input parameter variation and the contour is either C1 or C3 in order to provoke porosity inside of the samples with a the influence of the energy input into the contour is evaluated. A second sample series consists of a filling only laser power, scan speed and the beam offset are modified but also the sequence of the exposure.This way parameters are varied in five different ways (C1-C5), whereas the filling parameters (F0) are kept constant. Not Throughout all experiments the layer thickness was kept constant at 120 μm.

Not Throughout all experiments the layer thickness was kept constant at 120 μm. Within the first sample series contour on an EOSINT P396 to build z-direction tensile specimens type 1A (DIN EN 527) for the first two sample series At the DMRC, Paderborn University the material VESTOSINT® Z2837(refresh rate 50/50) is processedhigher mechanical properties. VESTOSINT® Z2837 showing higher recyclability and VESTOSINT® 3D 8754 HT1 showing the presented x-build direction applying adequate process parameters. In this work it is used in two different formulations, As presented in Table 1 the material VESTOSINT® 3D 8754 HT1 can reach an elongation at break of 40 % in properties and slightly higher melt temperature developed by Evonik particularly for the laser sintering process.

For the first sample series the focus was the variation of contour exposure parameters to evaluate their effect on mechanical properties. Figure 2 shows tensile strength, elongation at break and the contour porosity for the different parameters. While quite low contour porosity of C2F0 and C3F0 results in high mechanical values, a very high contour porosity detected for sample numbers C5F0 and C4F0 causes very poor elongation at break. On the other hand, omitting extra contour exposure completely (F0) leads to acceptable tensile properties. The high contour porosity of C5F0 has its origin in the quite high energy input of the contour parameters of C5 which leads to a porous ring and therefore to premature material failure. The porous ring of C5F0 gets visible in Figure 3 and is already discussed in [3].

The porous ring of C5F0 gets visible in Figure 3 and is already discussed in [3]. Also, C4F0 shows more pores in the contour region in comparison to example C1F0.60500CIF0C2F0C3F0C4F0CSFOF0■ Tensile strength■ Elongation at break■ Contour porosityFigure 2: Tensile properties and contour porosity of the first sample row739Figure 3: XCT-scan pictures of tensile specimens cross sections of the 1st sample rowWithin the second sample series the filling exposure parameters are varied, so that the porosity inside of the sample is depicted together with the mechanical properties in Figure 5. High porosity could be generated for example inside of the samples C1F2 and C3F2. The contour exposure parameters C1 and C3 which led in sample row 1 to an improvement of the mechanical characteristic generate actually a dense contour region which is visible in Figure 4 but could not compensate the high number of defects inside of the samples.

Nevertheless, samples built with contour parameters C3 on average show higher mechanical properties, while the filling porosity is not clearly lower in comparison to the ones of C1. Comparison of C3F6 with F0 shows quite similar elongation at break and tensile strength although the porosity of F0 is higher. C3F5 results as well in a high elongation at break while the filling porosity is comparatively poor. Here, neither contour exposure nor porosity seem to be the determining factors.Figure 4: XCT-scan pictures of tensile specimen cross sections of the 2nd sample row740Furthermore, contour and filling exposure seem mutually interact concerning the total energy input. Regarding C3F5 and C3F6 contour exposure parameters are indeed the same but pictures in Figure 4 show that C3F5 has more big pores in the core and furthermore that the cross section is enlarged.

Regarding C3F5 and C3F6 contour exposure parameters are indeed the same but pictures in Figure 4 show that C3F5 has more big pores in the core and furthermore that the cross section is enlarged. The energy input by F5 might be that high, that material is melted on beyond the contour region and the observed pore development inside of the sample might be accounted to material degradation. Nevertheless, the contour is quite dense with a porosity of 3.57 % and mechanical characteristics are still not decreased. On the contrary C1F4 shows only slightly higher filling porosity but a more porous contour region which results in reduced mechanical properties. Viewing the XCT-pictures in Figure 4 reveals further that the shape of the pores are different from the ones of C3F5. Here incomplete particle coalescence due to insufficient laser energy input seems to be the reason for the pore development.

Here incomplete particle coalescence due to insufficient laser energy input seems to be the reason for the pore development. With increasing total energy input the coalescence of polymer particles is favored and therefore porosity reduced.6050ro~ 40 ---- ,..c; to 5 30 b r./lCl)~ 20 i::: ~0ClFlClF2ClF3ClF4C3FlC3F2C3F5C3F6FO■ Tensile st1ength■ Elongation at break■ Filling porosityFigure 5: Tensile properties and filling porosity of the second sample rowThe porosity of both sample series are directly reflected in measured part densities. In Figure 6 the filling porosity evaluated from the XCT-scans and sample density which was measured by immersed body method is depicted for both sample rows. As to be expected, density is low when the filling porosity is high and vice versa.

However, if a given density around 1 g/cm³ is reached, mechanical properties are only weakly sensitive to porosity / density any more. The figure shows not only densities and mechanical properties of samples from the DMRC but also of the ones manufactured and tested at Evonik with a different material VESTOSINT® 3D 8754 HT1 which is not recyclable. The quite low tensile strength at a high density of 1.038 g/cm³ is the result of the high contour porosity of C5F0. However, the strong scattering of mechanical properties even for similar filling porosity indicate, that there is origin, most probably in other microstructural properties.603000Cl:l 50 ~ -..... 40 t~ 30 .b r/1■ ■X■X++ "\.+■ ' ...+X ♦ ■•• X@X ■2500Cl:l2000 ~ .......

While injection molded parts show small, equally distributed crystals in the middle and a less crystalline, more amorphous shell for both materials (PA12 and PA613), big differences in laser sintered parts are visible. For PA12, the difference in crystallinity is expected and known, but PA613 shows, besides typical layers and bubbles of laser sintering processes, also amorphous sections. This is a significant difference to the high crystallinity and “big” spherulites of PA12. Some of the parts made out of PA613 show significant higher values with respect to elongation at break. Analysed tensile bars with outstanding 80 % elongation at break show less spherulites than tensile bars that reached 30 % elongation.

Analysed tensile bars with outstanding 80 % elongation at break show less spherulites than tensile bars that reached 30 % elongation. The study is continued and so far, there is no proof that crystal structure and properties correlate.Injection molded specimensLaser sintered specimenslniection molded specimens: Both, PA12 and PA6x show small, equally distributed crystals and a less crystalline, more amorphous shell.SLS specimens: PA12 shows high crystallinity and "big" spherulites, PA6x also amorphous sectionsFigure 9: Light spectroscopy images of injection molded and laser sintering specimens743Conclusion and outlookIn this work, different laser parameter variations of the contour and filling strategy have been performed with to different materials at the DMRC and at Evonik to investigate the correlation between PA613 specimen’s microstructure and corresponding mechanical properties.

Besides porosity analysis, density and crystallinity have been tested to find the domaining reason for component failure. It becomes obvious, that density measurement reproduces porosity results of XCT-scans pretty well. Further, it could be shown that provoked porosity inside of the sample and a porous ring in the contour region as well lead to reduced tensile properties. Nevertheless, when a certain degree of density is reached, neither porosity nor crystallinity seem to be decisive. As the influence of crystallinity might be overlapped by the high provoked porosity and first considerations of spherulites were promising further investigations are planned. By a variation of exposure parameters resulting all in dense samples the effect of crystallinity shall be evaluated whereas porosity is avoided in the future.

By a variation of exposure parameters resulting all in dense samples the effect of crystallinity shall be evaluated whereas porosity is avoided in the future. Further origins of fluctuations especially of elongation at break might be crystallite size, polymer chain length and surface defects which will also be regarded.Overall, the PA6 like material is more sensitive to process parameters in comparison to standard LSmaterial PA12 but can get also closer to injection molded PA6 properties when using suitable settings.References [1] Kantesh Balani, Vivek Verma, Arvind Agarwal, and Roger Narayan. Physical, Thermal, and MechanicalProperties of Polymers. Biosurfaces 2015.[2] C. E. Majewski, H. Zarringhalam, N. Hopkinson. Effects of Degree of Particle Melt and crystallinity inSLS Nylon-12 parts.[3] C. Kummert and H.-J. Schmid. The Influence of Contour Scanning Parameters and Strategy on Selective[4]Laser Sintering PA613 Build Part Properties.

fluence of Post Curing Conditions on the Mechanical Properties of Stereolithographic PhotopolymersSuresh Jayanthi, Bronson Hokuj and John Lawton DuPont Somos™ Materials Group Two Penn's Way, Suite 401 New Castle, DE 19720 Tel:302-328-5435; Fax:302-328-5693Post cure methods have always been a critical issue for most of the currently used stereolithography photopolymers. Different post cure conditions can be used to enhance the mechanical and physical properties of the stereolithography models. To better understand the influence of post cure conditions on these material properties, an experimental study evaluating various post cure conditions has been conducted. Ultra-violet, thermal and ultra-sound post curing techniques are investigated and a comparative study is made for one epoxy based photopolymer - Somos™ 6100.

Ultra-violet, thermal and ultra-sound post curing techniques are investigated and a comparative study is made for one epoxy based photopolymer - Somos™ 6100. Similar results should be attainable for other epoxy based photopolymers.The post curing stage still remains a critical element in the processing of models built using stereolithography. The eventual mechanical properties as well as the physical properties are greatly influenced by the post curing conditions used. Thermal and the ultra-violet post curing techniques have been in use for quite some time now. In this study a third approach involving ultra-sound post cure has been investigated. Initial results show that the ultra-sound post cure is better than the UV post cure. However, since the physical properties are very much dependent on the photopolymer material, no generalizations to other materials can be made. The photopolymer used in this study is an epoxy based resin called Somos™ 6100.

The photopolymer used in this study is an epoxy based resin called Somos™ 6100. Epoxy based photopolymers in general have substantially better mechanical properties than acrylate photopolymers. Further improvements in the properties of the epoxy based green parts is achievable by a proper choice of the post curing conditions.To a large extent the most common post cure techniques employed in curing green parts are UV and thermal cure. UV cure is intended to directly polymerize any uncured liquid trapped with the solid by irradiating the part with broad-band ultra-violet light. Parts subjected to additional UV exposure show an increased concentration of acid initiators resulting in additional polymerization. On the other hand thermal cure induces additional polymerization in the green part by increasing the mobility of the polymer chains through higher temperatures.

On the other hand thermal cure induces additional polymerization in the green part by increasing the mobility of the polymer chains through higher temperatures. Yet another approach which can induce additional polymerization in a green part is ultra-sound curing where the degree of cross-linking is increased through a combination of vibrational and some thermal post cure. The polymerized chains of epoxy molecules within the object matrix still contain some active groups and movement of the chain segments induced by ultra-sound vibrations and rise in temperature bring together unreacted groups causing additional monomer conversion and107cross-linking. In general, in the epoxy based materials, the photopolymer cure is initiated by the laser exposure. This laser curing will theoretically continue in time until the glass transition temperature(Tg) is reached.

This laser curing will theoretically continue in time until the glass transition temperature(Tg) is reached. Above the glass transition temperature there is enough mobility within the object matrix to allow the acid molecules to initiate polymerization and for the unreacted groups to find each other and cross-link. Below the glass transition temperature, however, the unreacted groups may not be able to cross-link due to lack of matrix mobility. Post cure techniques that elevate the object temperature above the Tg should produce better mechanical properties. Table 1 presents a summary of the average values of v~rious mechanical properties under different curing conditions.

The number of samples in each set was five. The UV post cure was done in a 3D systems, Inc. PCA-250 using 10 Philips TLK 40W / 05 bulbs. The thermal post cure was done using a BLUE-M programmable convection oven. In the ultra-sound post curing, a standard test condition was used throughout the study.108HDT (C)59.669.1*67.560.351.252.861.349.566.859.5All the test samples in the ultra-sound post cure were subjected to the vibrational cure (using a Branson™ 220 ultra-sound cleaner operating at a frequency of 25 kHz) in a bath of silicone oil held at 80°C. All sample parts were post cured one hour following the completion of the part build. Parts were built without supports on a sheet of Mylar® glued to the platform, thus minimizing post processing. They were imaged by an Ar+ laser operating at 351 +364 nm using an overlapped X-hatch with a 2 mil hatch spacing and 5 mil beam diameter. The parts were tested within one week of post curing.

They were imaged by an Ar+ laser operating at 351 +364 nm using an overlapped X-hatch with a 2 mil hatch spacing and 5 mil beam diameter. The parts were tested within one week of post curing. ASTM procedures were followed for all tests.Discussion of Material Properties:One of the initial results from this study is that the different post curing approaches, namely photo, thermal and mechanical post curing, have their relative merits and demerits in inducing better mechanical properties to the resulting parts. Though this study cannot claim to be comprehensive it nevertheless provides some valuable insights into understanding the influence of different post curing conditions for an epoxy based photopolymer. The following discussion explains the different mechanical properties and compares the relative merits of the various post curing methods studied.

The following discussion explains the different mechanical properties and compares the relative merits of the various post curing methods studied. The different sets of post curing conditions represent typical ranges of thermal and UV post cure conditions generally used in the industry. It is observed that different curing conditions affect the various mechanical properties in differently. Table 2 lists the material property and the best and worst curing conditions based on the maximum and minimum values obtained as illustrated in the accompanying charts.# Material Property1Tensile Modulus2Tensile Strength3Tensile Elongation4Flexural Strength5Flexural Modulus6Izod Strength7Heat Deflection8Shore HardnessHighest Max. Value Meas.

It is observed that the ultra-sound curing in combination with thermal109curing offers improved mechanical properties. In general, a combination of cure techniques is preferred over a single technique. The most common technique, 60 min UV post cure, frequently produces the lowest maximum properties.Tensile Properties:In the evaluation of tensile properties, the three properties of interest are the modulus of elasticity, tensile strength and the percent elongation. Stress-strain characteristics for plastics almost always show a linear region at low stresses, and a straight line drawn tangent to this portion ofthe curve permits calculation ofan elastic modulus. 1 This modulus is defined as the ratio of stress to the corresponding strain below the proportionality limit of a material.

1 This modulus is defined as the ratio of stress to the corresponding strain below the proportionality limit of a material. The modulus of elasticity ( tensile modulus ) is calculated by extending the linear portion of the load-extension curve and dividing the difference in stress corresponding to any segment on this straight line by the corresponding difference in strain. Many polymeric materials are visco elastic in their stress-strain behavior and display significant non-linearity and strain-rate sensitivity. For materials where no proportionality is evident, the secant modulus is calculated.

For materials where no proportionality is evident, the secant modulus is calculated. Secant modulus is calculated at a designated strain (typically 1% or 5%) and is obtained by dividing the corresponding stress by the designated strain.30 min @ 100 C+30 minUV30 min UV+30 min@100 C30 min ultra-sound+30 minUV30 min ultra-sound+30min@130 C 30 min ultra-sound+30min@SOC60 min UV cure60 min ultra-sound insilicone oil@SO C60 min @ 150 C60 min @ 130 C60 min @ SO C[;J rangeo10002000300040005000[;J minTensile Modulus in MPaFigure 1: Tensile Modulus per ASTM D 638All test results reported here were measured at pull rates of 0.2 in/min. Specimens were type Ml with a 1 cm square cross-section. Figure 1 shows the ranges of values for the variousI ASTM D 638 - Standard Test Method for Tensile Properties of Plastics. Annual Book of ASTM Standards 1993. Section 8.110post cure conditions.

Figure 1 shows the ranges of values for the variousI ASTM D 638 - Standard Test Method for Tensile Properties of Plastics. Annual Book of ASTM Standards 1993. Section 8.110post cure conditions. The ultra-sound approach seems to provide improved modulus values because of the higher mobility induced into the polymeric chains which allows the various unreacted chains to move closer and join together. The induced mobility followed by additional polymerization in a high temperature medium makes the material stronger. Another interesting result observed from the table below is that thermal cure followed by the UV cure yields better results as opposed to UV cure followed by the thermal cure under similar circumstances.

This relationship seemed to hold for several other properties as welL30 min @ 100 C+30 min UV30 min UV+30 min@100 C30 min ultra-sound+30 minUV30 min ultra-sound+30min@130 C30 min ultra-sound+30min@SO C60 min UV cure60 min ultra-sound in siliconeoil@SO C60 min @ 150 C60 min @ 130 C60 min @ SO Crangeo1 0203040506070minTensile Strength in MpaFigure 2: Tensile Strength per ASTM D 638The tensile strength is the maximum tensile stress sustained by the specimen during a tension test. It is calculated by dividing the maximum load by the cross-sectional area of the specimen. Not much variation is observed in the tensile strength values. Though considerable variation is observed in the modulus, the small variations in the tensile strength of the samples suggests that the specimens are subjected to different levels of induced strain.

Though considerable variation is observed in the modulus, the small variations in the tensile strength of the samples suggests that the specimens are subjected to different levels of induced strain. Figure 2 shows that the ranges of variation of the tensile strengths in almost all the different post curing conditions are quite narrow suggesting that the material attains the optimal tensile strength during the building stage itself. Interestingly, the most widely used technique, UV only post cure, produced the lowest values.The elongation of the test specimen as a percent of the gage length is evaluated to determine elongation behavior of the specimen under load. It is obtained by dividing the changeill80in gage length at the moment the applicable load is reached, by the original length. The result is expressed as a percentage. The experimental testing reported here was done according to the ASTM D 638 standard. Figure 3 shows the percent elongation under different conditions.

The result is expressed as a percentage. The experimental testing reported here was done according to the ASTM D 638 standard. Figure 3 shows the percent elongation under different conditions. It is seen that the percent elongations are lower under ultra-sound curing combined with thermal cure. Highest elongation values are obtained by curing the parts at 1000e for 30 minutes followed by a 30 min UV cure.

Highest elongation values are obtained by curing the parts at 1000e for 30 minutes followed by a 30 min UV cure. In general it is observed that the vibrational cure lowers the elongation values while the thermal cure increases these values.30 min @ 100 C+30 min UV30 min UV+30 min@100 C30 min ultra-sound+30 minUV30 min ultra-sound+30min@130 C30 min ultra-sound+30min@80C60 min UV cure60 min ultra-sound in siliconeoil@80 C60 min @ 150 C60 min @ 130 C60 min @ 80 Crangeo51 01 520minElongation 0/0Figure 3: Tensile Elongation per ASTM D 638Flexural Properties:The two main flexural properties of interest in the evaluation of photopolymer materials are flexural modulus and flexural strength. These properties are evaluated using a three point loading system: center loading on a simply supported beam. When a simply supported beam of homogenous elastic material is tested in flexure the maximum stress is induced in the outer fibers at the midspan.

When a simply supported beam of homogenous elastic material is tested in flexure the maximum stress is induced in the outer fibers at the midspan. This stress is designated as the flexural strength of the material at the moment of break. In case the test piece does not fail, the corresponding flexural strength at yield is evaluated. For materials that do not break at the maximum fiber strain (5% for ASTM 790), there exists a point on the load-deflection curve at which an increase in the deflection112occurs without a corresponding increase in load. This point is used to calculate flexural strength at yield. In either case, whether in failure or in yield, the flexural strength is calculated.As shown in Figure 4, not much variation is observed in the flexural strength. Interestingly, parts post cured just under UV showed slightly lower values than parts cured under conditions involving UV cure as well as thermal cure or UV cure as well as ultra-sound cure.

Interestingly, parts post cured just under UV showed slightly lower values than parts cured under conditions involving UV cure as well as thermal cure or UV cure as well as ultra-sound cure. This suggests that the increased mobility in the polymer chains resulting from either the thermal or vibrational cure helps improve the flexural strength. Both thermal and vibrational curing approaches yield similar results.30 min @ 100 C+30 min UV30 min UV+30 min@100 C30 min ultra-sound+30 minUV30 min ultra-sound+30min@130 C30 min ultra-sound+30min@80 C60 min UV cure60 min ultra-sound in siliconeoil@80 C60 min @ 150 C60 min @ 130 C60 min @ 80 Crangeo51 01 520minFlexural Strength in MpaFigure 4: Flexural Strength per ASTM D 790Flexural modulus is the ratio of the stress to the corresponding strain within the elastic limit. In case no apparent proportionality is observed, the flexural secant modulus is calculated.

In case no apparent proportionality is observed, the flexural secant modulus is calculated. It is the slope of the straight line that joins the origin and a selected point on the actual stress strain curve. The experimental testing reported here was done according to the ASTM D 790 standard.Figure 5 shows the flexural modulus values for the ten different cure conditions. All curing conditions yielded more or less similar ranges of flexural properties, suggesting that the nature of113the post cure is not influential in varying these properties. Results for the parts cured at 100°C for 30 minutes followed by a 30 min UV cure are not available. In general, the thermal and vibrational cures do not significantly enhance the flexural strength of the samples.

In general, the thermal and vibrational cures do not significantly enhance the flexural strength of the samples. The slightly improved values obtained using the thermal and vibrational cures as opposed to just the UV radiation cure suggests that the polymer chain mobility which results due to the additional energy supplied using these techniques is responsible for inducing additional cure.30 min @ 100 C+30 min UV30 min UV+30 min@100 C30 min ultra-sound+30 minUV30 min ultra-sound+30min@130 C30 min ultra-sound+30min@SO C60 min UV cure60 min ultra-sound in siliconeoil@SO C60 min @ 150 C60 min @ 130 C60 min @ SO Crangeo100200300400500DminFlexural Modulus in MpaFigure 5: Flexural Modulus per ASTM D 790Hardness Properties:In this test a very hard indenter is pressed under a load into the surface of the materiaL For the photopolymer materials the most commonly used measure is shore hardness (Shore D).

Here the indenter is a steel pencil in the form of a rounded cone. With a prescribed force exerted by a spring, the pencil is pressed into the material and the indentation depth is measured on a scale ranging from 0 to 100. Shore (durometer) hardness is a measure of the indentation hardness of the material and is inversely proportional to the penetration. The indentation hardness is dependent on the elastic modulus and the visco-elastic behavior of the materiaL However, the property only allows an empirical estimation of the hardness since there exists no simple relationship between the shore hardness and any fundamental material property. Type D114Shore durometer is typically used for reporting the hardness of photopolymer materials. The experimental testing reported here was done according to the ASTM D 2240 standard. Figure 6 reports the shore D values. The maximum values of hardness are obtained under a 30 minute ultra-sound cure followed by a 30 minute thermal cure at 130°C.

The typical method for evaluating the impact strength is the Izod type test in which a notched specimen is held as a vertical cantilever beam and is broken by a single swing of a pendulum with the line of initial contact at a fixed distance from the specimen clamp. (Some vendors have reported results for un-notched specimens or have not specified whether the specimens were notched. This data should only be compared to the data for notched specimens.) The experimental testing reported here was carried out according to the ASTM 256 standard. 7 shows the impact strength results. The results from the samples cured for 30 minutes in a ultra-sound chamber followed by a thermal cure for 30 minutes in an oven at 130°C are not115available. In general, thermal cure gives better impact strength compared to vibrational or radiational cures.

In general, thermal cure gives better impact strength compared to vibrational or radiational cures. However, as can be seen from the figure below, considerable variations are observed within in each data set.30 min @ 100 C+30 min UV30 min UV+30 min@100 C30 min ultra-sound+30 minUV30 min ullra-sound+30min@130 C30 min ullra-sound+30min@80C60 min UV cure60 min ultra-sound in siliconeoil@80C60 min @ 150 C60 min @ 130 C60 min @ 80 Crangeo0.20.40.60.81.2DminIzod Strength in Ft-Ib/inFigure 7: Izod Impact (Notched) Strength per ASTM D 256Heat Deflection Temperature:This test provides for the determination of the temperature of the material at which a certain deformation is induced under flexural load of the test specimen. The test provides a measure of the thermo-mechanical stability of the material.

The test provides a measure of the thermo-mechanical stability of the material. A rectangular specimen is tested as a simple beam while immersed in a heat transfer medium with a load applied at its center to give a maximum fiber stress of 66 psi. Higher loads were not used because there is no advantage to using higher loads when measuring deflection temperature of the present-day plastics with present-day instruments2. With the bar immersed and the load applied, the temperature is raised at 2° C/min. The temperature at which a deflection of 0.010 inches occurs is noted as the heat deflection temperature under flexural load of the material. The experimental testing reported2ASTM D 648 - 82 (Reapproved 1988) - Standard Test Method for Deflection Temperature of Plastics Under Flexural Load. Annual Book of ASTM Standards 1993. Section 8.116here was done according to the ASTM D 648 standard.

Annual Book of ASTM Standards 1993. Section 8.116here was done according to the ASTM D 648 standard. Figure 8 shows the results of the heat deflection study under different post curing conditions. Thermal cure is more effective than the vibrational or radiational cures . Thermally induced distortions due to the effects of gravity in the samples can be minimized by placing the samples in a dense medium like silicone oil.30 min @ 100 C+30 min UV30 min UV+30 min@100 C30 min ultra-sound+30 minUV30 min ultra-sound+30min@130 C30 min ultra-sound+30min@80 C60 min UV cure60 min ultra-sound in siliconeoil@80 C60 min @ 150 C60 min @ 130 C60 min @ 80 Crangeo20406080DminDeflection Temperature in CFigure 8: Heat Deflection Temperature per ASTM D 648Acknowledf:ments:This presentation is a result of the efforts of many people in the Somos™ Materials Group.

We would like to acknowledge the valuable assistance rendered to us by Harry Bacon and Al Colombo in fabrication and evaluation of the test-pieces. We would like to thank Dr. Daniel Mickish and Dr. Glen Thommes for their continuous support and advice.Inquiries:Inquiries regarding this study may be directed to Suresh Jayanthi at the address listed or bye-mail tojayantss@esvax.dnet.dupont.com.117.

E INFLUENCE OF POST-SLS-BUILD ANNEALING ON NYLON 11 MATERIAL PROPERTIESLauren A. Slattery, Nicole L. Guckert, Courtney E. Shell, and Richard R. NeptuneDepartment of Mechanical Engineering, The University of Texas at Austin, Austin, TX 78712AbstractFunctional energy storage and return prosthetic and orthotic devices have been manufactured out of Nylon 11 using selective laser sintering due to its high ductility and energy return properties. However, there is concern that material voids caused by incomplete sintering may compromise material properties and lead to premature fracture. Post-build annealing has the potential to eliminate voids caused by incomplete sintering and increase part ductility and strength.

Post-build annealing has the potential to eliminate voids caused by incomplete sintering and increase part ductility and strength. The purpose of this study was to post-build anneal Nylon 11 tensile specimens at 1) slightly below their melting temperature, 2) their recrystallization temperature, and 3) their glass transition temperature for two different time durations (12 and 24 hours) to assess the effectiveness of annealing in improving ductility and strength. Specimens annealed at their glass transition temperature had significantly greater percent elongations and lower Young’s moduli than specimens annealed close to their melting or recrystallization temperatures. At each temperature, specimens annealed for shorter durations demonstrated a greater increase in percent elongations and a greater decrease in Young’s moduli.

At each temperature, specimens annealed for shorter durations demonstrated a greater increase in percent elongations and a greater decrease in Young’s moduli. Annealing at the glass transition temperature for 12 hours resulted in the highest percent elongation, although it was not significantly different from the control (unannealed) specimens. However, at these annealing conditions Young’s modulus significantly decreased from the control specimens. Across all annealing conditions, Young’s modulus and percent elongation were found to be negatively correlated. Future work should focus on annealing specimens for additional combinations of temperature and duration to further improve ductility while minimizing the negative effects on part strength.IntroductionEnergy storage and return (ESAR) prosthetic devices are commonly manufactured out of carbon fiber, which requires extensive production costs and time.

In contrast, selective laser sintering (SLS), an additive manufacturing technique, has proved to be a powerful alternative involving less production time and cost while allowing for more complex, customizable designs. Nylon 11 has been used to manufacture ESAR prosthetic and orthotic devices due to the material’s desirable properties, including high flexural strength and ductility and low energy dissipation [1-3]. However, there is concern that small voids caused by incomplete sintering may compromise material properties and lead to premature fracture.Post-build annealing of SLS parts has the potential to decrease the number of voids and improve material properties through refining the material structure and making it more homogeneous. When the temperature of Nylon 11 is incrementally increased, it passes through three phases, with the transition points between these phases being the glass transition565temperature, recrystallization temperature and melting temperature [4].

Annealing just below the melting temperature (~186°C for Nylon 11) may complete the sintering and eliminate remaining voids [5]. However, annealing at high temperatures has also been shown to decrease the ultimate elongation [6]. Alternatively, annealing in the recrystallization range may allow the material to form more perfect crystals, improve the regularity of polymer chains, and increase interlamellar ties [4]. The increase in interlamellar ties reduces brittleness and increases ductility [4]. Studies have shown that Nylon 11 crystallinity reaches a maximum when annealed at 165°C [7]. Annealing between the glass transition temperature (Tg = 45°C) and recrystallization temperature can reduce residual stresses and increases movability of chains [7]. Within this range, annealing at a temperature closer to the glass transition temperature has been shown to improve ductility [6].

Within this range, annealing at a temperature closer to the glass transition temperature has been shown to improve ductility [6]. However, improving ductility through annealing can adversely affect material strength [6].The purpose of this study was to examine the influence of post-SLS-build annealing on Nylon 11 material properties to identify the combination of annealing temperature and duration that maximizes part ductility without adversely affecting part strength. Specifically, we analyzed Nylon 11 tensile specimens annealed at 1) slightly below their melting temperature, 2) their recrystallization temperature, and 3) their glass transition temperature for two different time durations (12 hours and 24 hours).

Based on previous work [6, 7], we expected that annealing between the glass transition and recrystallization temperatures would yield the greatest improvement in ductility with only modest changes in strength.MethodsNylon 11 tensile specimens (n=154) were manufactured using SLS, annealed at 80°C, 165°C or 186°C for 12 or 24 hours, and mechanically tested to determine specimen ductility and strength quantified by percent elongation and Young’s modulus, respectively.Nylon 11 Test SpecimensSpecimens were manufactured using a blended 50/50 mix of virgin and single-use overflow Nylon 11 powder (PA D80-ST, Advanced Laser Materials, Temple, TX) in a Vanguard HiQ/HS Sinterstation (3D Systems Corporation, Rock Hill, SC) using the temperature parameters specified in Table 1. A 14 x 11 array of 154 tensile specimens designed using ASTM D638 was created.

A 14 x 11 array of 154 tensile specimens designed using ASTM D638 was created. Specimens were placed in the build volume such that the longest dimension of the specimen was parallel to the movement of the roller, with each specimen 0.15” apart in both the horizontal and vertical directions (Fig. 1). A heat shield was added on the top and bottom of the build volume to facilitate more gradual cooling of the specimens.

Once the build was complete and allowed to cool, the specimens were separated from the unsintered powder, bead blasted and air blown to remove excess powder.566Warm-UpBuildCooldownTime0:00:00 0:08:58 2:00:02 2:48:26 2:48:26 12:59:11 12:59:11 14:01:44 17:40:13 43:18:28Part Bed34 100 188 188 188 188 188 185 57 35Temperature [°C] Right Feed Bin 33 118 142 144 144 141 141 92 49 34Left Feed Bin 33 97 141 142 142 142 142 90 44 33Cylinder27 138 138 138 138 138 138 85 46 28Piston27 81 150 150 150 150 150 115 66 27Table 1: Temperature parameters of the SLS build for a blended 50/50 mix of virgin and single-use overflow Nylon 11 powder in a Vanguard HiQ/HS Sinterstation.RollerFigure 1: Schematic of the SLS build volume containing a 14x11 array of tensile specimens and a heat shield on both the top and bottom of the build.AnnealingSpecimens were then annealed at three different temperatures (80°C, 165°C or 186°C) for two different durations (12 or 24 hours) yielding six annealing temperature-duration combinations.

In addition, a control set that received no annealing was used for comparison. To control for variations in material properties across the build, specimens in the four outermost columns, two on each side, of the 14x11 array were randomly assigned to condition sets and specimens in the central seven columns were systematically distributed to the seven condition567sets. A temperature-controlled oven (Model 5851, National Appliance Company, Portland, OR) was used to anneal each group of test specimens. For each condition set, specimens were placed on a rack 1” above the oven’s hot plate. The rack was used to render a more uniform temperature distribution throughout the test specimen than direct placement on the heating element. To prevent oxidation, the door was sealed and the oven chamber was evacuated and backfilled with nitrogen, which minimized the amount of oxygen remaining in the oven.

To prevent oxidation, the door was sealed and the oven chamber was evacuated and backfilled with nitrogen, which minimized the amount of oxygen remaining in the oven. At the end of the desired annealing duration, the heating element was turned off and the specimens were allowed to cool slowly until the oven reached room temperature.Tensile TestingThe material properties of the tensile specimens were determined using ASTM D638 (Instron 3345, Norwood, MA). An extensometer was used to improve resolution until the strain reached a value of 0.055 mm/mm, which corresponded to a point after the end of the elastic region of the stress-strain curve. From these measurements, percent elongation and Young’s modulus were determined and compared across annealing conditions (temperature, time).Statistical Analysis:A two-way (temperature, time) MANOVA was used to compare the percent elongation and Young’s modulus across conditions using SPSS 16.0 (SPSS, Inc. Chicago, IL).

Chicago, IL). When significant differences were detected, Tukey pair-wise comparisons were performed to determine which condition sets were significantly different from the others and identify the condition set that most improved percent elongation. Independent sample t-tests were then used to compare the percent elongation and Young’s modulus in this condition set to those in the control (unannealed) set. The equal variance assumption was tested by Levene’s Test for Equality of Variances for both independent sample t-tests to ensure the appropriate t-test was performed.

The equal variance assumption was tested by Levene’s Test for Equality of Variances for both independent sample t-tests to ensure the appropriate t-test was performed. The Pearson’s correlation coefficient was calculated to determine the relationship between percent elongation and Young’s modulus.ResultsWithin the annealing condition sets, percent elongation significantly increased and Young’s modulus significantly decreased as annealing duration and temperature decreased (all p≤0.000), but the interaction between duration and temperature did not influence the percent elongation (p=0.200) or Young’s modulus (p=0.4100). Annealing at 80°C increased the percent elongation compared to annealing at 186°C (p=0.000) and 165°C (p=0.000), but there was no difference between the percent elongation of specimens annealed at 186°C and 165°C (p=0.138) (Fig. 2).

2). Annealing at 80°C decreased the Young’s modulus compared to annealing at 186°C (p=0.000) and 165°C (p=0.000), while annealing at 165°C also significantly decreased the Young’s modulus compared to the specimens annealed at 186°C (p=0.029) (Fig. 3). At each temperature, annealing for the shorter duration (12 hours) increased percent elongation (p=0.000) and decreased Young’s modulus (p=0.000).The specimens annealed at 80°C for 12 hours produced the highest average percent elongation of all conditions, but they had the lowest average Young’s modulus. Although the highest average percent elongation occurred in this annealing condition set, the increase in percent elongation was not significant when compared to the control set (p=0.082) while the568decrease in Young’s modulus was significant compared to the control set (p=0.001).

A negative correlation (p=0.000, R=-0.701) was identified between the Young’s moduli and percent elongations, indicating that post-build annealing of SLS tensile specimens had opposite effects on material ductility and strength.Figure 2: Average percent elongation and standard deviation (SD) of annealed Nylon 11 specimens. Annealing duration had a significant effect on percent elongation at each temperature. Significant differences between annealing temperatures are indicated with an asterisk (*).569Figure 3: Average Young’s modulus and standard deviation (SD) of annealed Nylon 11 specimens. Annealing duration had a significant effect on Young’s modulus at each temperature. Significant differences between annealing temperatures are indicated with an asterisk (*).DiscussionAs specimens were annealed for shorter durations at lower temperatures, percent elongation increased while Young’s modulus decreased.

This trend, in addition to the negative correlation identified between the two variables, indicated that annealing for a shorter duration might further improve ductility but at the expense of decreased strength. Since percent elongation increased at the shorter time duration, a post-hoc test was performed to assess whether even shorter time durations would further improve the ductility. In this post-hoc test specimens were annealed at 80°C for 6 hours. There was no significant difference in percent elongation (Fig. 4) or Young’s modulus (Fig. 5) when annealed for 6 hours versus 12 hours (p=0.225 and p=0.257, respectively). However, there was an increase in percent elongation (Fig. 4) and decrease in Young’s modulus (Fig. 5) relative to the control (unannealed) specimens (p=0.009 and p=0.000, respectively).570Figure 4: Average percent elongation and standard deviation (SD) of annealed Nylon 11 specimens from the post-hoc trials and previous condition sets.

Significant differences between annealing conditions are indicated with an asterisk (*).Figure 5: Average Young’s modulus and standard deviation (SD) of annealed Nylon 11 specimens from the post-hoc trials and previous condition sets. Significant differences between annealing conditions are indicated with an asterisk (*).These results suggest that the annealing parameters that significantly improve percent elongation compromise material strength, which is consistent with previous work with nylon rod stock [6]. The Young’s moduli from the specimens annealed at 80°C for 12 and 6 hours (Fig. 5) fall outside the range of acceptable values (1400-1800 MPa). Thus, although ductility was improved, material strength was compromised.

5) fall outside the range of acceptable values (1400-1800 MPa). Thus, although ductility was improved, material strength was compromised. Future work should include exploration of571additional annealing temperature and duration combinations to improve ductility while minimizing the negative effects on strength.One potential limitation of this study is that during the annealing process, the specimens were placed on a rack one inch above the oven hot plate, potentially resulting in uneven heating of the specimens. However, since the annealed specimens were less than 4 mm thick, the temperature distribution within each specimen was likely uniform. However, this annealing configuration may not be ideal for thicker parts. Thus, future work should include development of a three- dimensional thermal model of parts during annealing to analyze the temperature distributions across parts with various thicknesses.

Thus, future work should include development of a three- dimensional thermal model of parts during annealing to analyze the temperature distributions across parts with various thicknesses. Such a model could be used to optimize the annealing configuration to achieve a uniform temperature distribution across specimens.ConclusionsThis study explored the influence of post-SLS-build annealing on Nylon 11 material properties. The results indicate that annealing at 80°C for 12 hours produces higher percent elongations and lower Young’s moduli than annealing at 165°C or 186°C. However, no significant difference in percent elongation was found between specimens annealed at 80°C for 12 hours and the control (unannealed) specimens. In addition, Young’s modulus was significantly decreased in those specimens annealed at 80°C for 12 hours compared to the control specimens.

In addition, Young’s modulus was significantly decreased in those specimens annealed at 80°C for 12 hours compared to the control specimens. Young’s modulus and percent elongation were found to be negatively correlated and post-hoc trials indicated that annealing at 80°C for a shorter duration further improved percent elongation, although it lead to further decreases in the Young’s modulus. Future work should focus on annealing specimens for additional combinations of temperature and duration to further improve ductility while minimizing the negative effects on material strength.AcknowledgmentsThis work was supported by VA grant 1 I01 RX000311 and the NSF Graduate Research Fellowship Program. The contents are solely the responsibility of the authors and do not necessarily represent the official views of the VA or NSF.572References1. Faustini M.C., Neptune R.R., Crawford R.H., Stanhope S.J.

Particle size distribution and powder layer inhomogeneity after layer coating influence melt pool dynamics and may cause defects in the work pieces. This correlation can hardly be analyzed in experiments, so a three-dimensional transient numerical simulation model is used. The model is based on the continuity equation, the heat equation and the Navier-Stokes equation. Therefore, the finite-volume method capabilities in OpenFOAM are used. The free surfaces of the multi-phase system are calculated using the volume of fluid method. The powder beds have unimodal or bimodal distributions without random effects in the particle composition. Their density and thermal conductivity is adapted to reality.

The powder beds have unimodal or bimodal distributions without random effects in the particle composition. Their density and thermal conductivity is adapted to reality. The investigations of the melt pool and the porosity formation demonstrate an advantage of more and smaller particles in the powder for compensation of defects in the powder bed, similar to the results of the experiments.IntroductionLaser beam melting (LBM), also named selective laser melting (SLM), is an additive manufacturing process, which is one of the most innovative manufacturing technologies. In the LBM technique a laser beam heats and melts lines of a material powder bed. The single molten tracks form a flat structure, which can have filigree as well as massive areas. After a new powder coating the process is repeated and the single layers compose a three-dimensional part.With this new process many materials like metals, ceramics, etc. can be treated.

After a new powder coating the process is repeated and the single layers compose a three-dimensional part.With this new process many materials like metals, ceramics, etc. can be treated. The workability is influenced by various consolidation mechanisms, as shown by current research [Kruth et al., 2007]. Therefore, the choice of the process parameter is crucial to get high quality products and to minimize residual porosity.In the current studies of this young manufacturing method the influence of individual process parameters is not yet understood sufficiently [Zheng et al. 2012]. In LBM various dynamical phenomena like wetting, solidification and grain growth are induced as a result of the high temperature gradients [Das, 2003]. To reduce expensive experiments and to enable direct observation of the process dynamics for analysis, numerical models are necessary.

To reduce expensive experiments and to enable direct observation of the process dynamics for analysis, numerical models are necessary. There are several simulation methods, which interpret the powder bed as homogeneous bulk material with adapted properties. A finite element method with thermo-mechanical effects was used toinvestigate the impact of heat on residual stresses and deformations [Zäh et al., 2009]. A first inhomogeneous powder layer was composed of regions with differing powder densities for a non-uniform behavior of the powder bed in the process [Xiao and Zhang, 2007]. The absorption of the laser energy and thermal conductivity in an irregular powder bed was determined using an analytical model [Gusarov and Smurov, 2010].

The absorption of the laser energy and thermal conductivity in an irregular powder bed was determined using an analytical model [Gusarov and Smurov, 2010]. A two-dimensional lattice Boltzmann method shows a significant effect of the powder density of the powder bed on the melt pool dynamics [Körner et al., 2011].All of these simulations do not show the influence of powder beds of various particle sizes and defects in the powder beds on the process. Since pores can arbitrarily appear in the workpiece, our approach is that defects in the powder layer after coating are potential seeds for these pores. However, these defects can occur in reality during layer coating (figure 1). Our second approach is that the impact of these coating defects can be affected by the size distribution of the powder particles. Therefore, our goal is to investigate the influence of inhomogeneities in the powder bed onto the melt pool characteristics and the process stability.

Therefore, our goal is to investigate the influence of inhomogeneities in the powder bed onto the melt pool characteristics and the process stability. Additionally, we want to show the effects of different particle size distributions on this investigation of the process behavior. By these findings, a suitable particle size distribution can reduce the porosity of the work piece and the part quality can be improved at usually too high scan speeds. So the productivity of the LBM process can be increased. Experiments have randomly distributed defects and an analysis of the process and the dynamics of pore development is hardly possible. In contrast, effects of coating defects and different particle size distributions on the process can be analyzed separately and visualized by numerical methods. This simulative analysis is conducted using our three-dimensional numerical simulation model for LBM. Finally, the results are compared to experimental data.GrooveFig. 1.

This simulative analysis is conducted using our three-dimensional numerical simulation model for LBM. Finally, the results are compared to experimental data.GrooveFig. 1. Groove in the powder bed after layer coatingSimulation ModelOpenFOAM (Open Field Operation And Manipulation, ©OpenCFD Limited) is a computational fluid dynamics toolbox. The software was developed to address continuum mechanical dynamical problems of heat transfer and multiphase flow using a finite volume formulation on an unstructured grid. The model for solving the specific physical phenomena of laser beam processing was developed at Institute of Photonic Technologies [Geiger et al., 2009]. This model was adapted for laser beam melting and capabilities for powder representation were implemented [Gürtler et al., 2013]. A combination of multiple physical effects is necessary for this process simulation.

A combination of multiple physical effects is necessary for this process simulation. The numerical scheme PIMPLE, a combination of PISO and SIMPLE, of OpenFOAM is used for the coupled calculation of the fluid dynamics which is given by the incompressible Navier-Stokes equationthe⃑⃑ ⃑ ⃑and the continuity equation for mass conservation⃑The volume of fluid (VOF) method describes the free surface between the molten metal and the ambient atmosphere. The surface dynamics is mainly determined by the evaporation pressure and surface tension.The phase transitions have the strongest influence on the temperature field. The thermal effects of fusion and evaporation are included in our model.

The thermal effects of fusion and evaporation are included in our model. The temperature field is calculated by iteratively solving the heat equation [Kohl et al., 2012]:⃑The laser beam is modeled as a continuum intensity field perpendicular to the powder bedwith Gaussian distribution and focal position at the top of the powder layer.The absorption on the material surface is calculated and the reflections result in discrete rays. They lead to additionally absorbed energy and end in the material at low ray intensity or after twenty reflections. The reflectance depends on the wavelength, angle of incidence and polarization, as specified by the Fresnel equations(√ √)for the parallel polarized part of the laser beam and(( √ ))for the vertical polarized part.

The material of the powder particles is AlSi12Mg (EN AC 43000) and it is produced under argon atmosphere by TLS Technik GmbH & Co. Spezialpulver KG.87% n i65t r a p e m u l o v4321002060 40 mean particle diameter in µm80100Fig. 2. The particle size distribution of AlSi10Mg produced under argon atmosphere:The measurement precision is 1.5 µm.The particle size distribution is measured with the dynamic image analysis method according to standard ISO 13322-2, see figure 2. This original powder is filtered in various size ranges. These fractions are mixed together by different volume fractions (table 3).

3. Illustration of the powder particles: a) SEM image, b) micrographic cross sectionIn our simulations, the size distribution between the upper and lower limit of a fraction is ̅̅̅ using the distribution of the particle sizes as averaging weightaveraged to one particle size (figure 2):̅̅̅∑ ∑with the particle size of each measurement point l, the normalized number of particlesand the mass of the particle This enables the representation of bimodal powder mixtures in the different blends.

Also they cannot be displayed as individual particles in the simulation volume except we drastically increase the spatial discretization. So the distribution size between 0 and 20 µm is reduced to 10 to 20 µm in the simulations.Table 5: Discrete particle size distributions adapted to simple ratiosPowder nameAl_20+32-40 Al_20+40-50 Al_20+50-63 Al_20+40_3-7 Al_20+40_4-6 Al_20+40_5-5Mean size of fraction 1 in µm 14.7 14.7 14.7 24.4 24.2 24.2Mean size of fraction 2 in µm 35.0 44.0 55.6 44.0 44.0 44.0Ratio of particles 3 : 2 3 : 1 6 : 1 5 : 2 4 : 1 6 : 1A regular arrangement of the powder particles is chosen to exclude random effects of the powder bed in the simulations and to have a plain test case. Therefore, we define unit cells of particles, which can easily be arranged for a powder layer or cluster. These unit cells have to include the correct particle quantities of both fractions.

Therefore, we define unit cells of particles, which can easily be arranged for a powder layer or cluster. These unit cells have to include the correct particle quantities of both fractions. Since we want to use only whole particles, the particle quantities of both fractions in the unit cells need to be integers. Thus, we have to slightly round the particle sizes of the fractions to reach small unit cells. In our simulation of single layers the unit cells have to be smaller than the height of the loose powder layer. The particle sizes and quantities of the different fractions in the unit cells that can be found in table 5 are used in the simulations.Furthermore the powder densities and the thermal conductivities of the powder beds haveto be considered.

The equation for the powder density is given aswith the volume of the unit cell (with height of simulated powder layer ), the ̅̅̅) and the̅̅̅̅single particle volume of each particle fraction quantity of each particle size in the unit cell.()(with mean sizeThe kind of arrangement of the powder particles is specified clearly. The powder density, however, can be modified by introducing and varying minimum distances between the spheres. Additionally, the distances reduce the interfaces between the particle spheres to decrease the thermal conductivity. Because of these two reasons, a distance of 2 µm is chosen to obtain the results shown in table 6 and 7. The resulting powder bed of the different particle size mixtures can be taken from figure 4.a) Al_20+32-40d) Al_20+40_3-7b) Al_20+40-50e) Al_20+40_4-6c) Al_20+50-63f) Al_20+40_5-5Fig. 4.

4. Illustration of the different powder beds (a-f): The connections between the single particles are caused by the displaying method: If there is no completely empty mesh cell (without metal) in between two particle spheres, the displaying algorithm cannot calculate a surface that separates the two particlesTable 6: Height and density of one powder layer in the simulationsPowder nameAl_20+32-40 Al_20+40-50 Al_20+50-63 Al_20+40_3-7 Al_20+40_4-6 Al_20+40_5-5Powder density of layer in sim. 0.47 0.49 0.45 0.48 0.49 0.47Height of sim. powder in µm 37.0 46.0 57.6 46.0 52.4 46.0Thickness of sim. layer in µm 17.4 22.5 25.9 22.1 25.7 21.6An investigation of the resulting powder layers delivers the results in table 6. The powder density of the powder layer varies between 45 % and 49 %. The height of the simulated powder layer mainly depends on the largest particles and ranges from 37 µm to 57.6 µm.

The powder standard DIN EN ISO 3923-1:2010-08 for loose powder. The values vary between 53 % and 58 %. These values cannot be compared to the powder density of the layers in the simulation directly (table 6), since the geometric boundary conditions differ. Thus, we have to determine the powder densities of simulated loose powder, which range from 45 % to 53 % (column 3 in table 7). Most powder mixtures agree with the real powders or are slightly undervalued. Additionally to this, the powder density of the powder layer in the simulation is less.

Most powder mixtures agree with the real powders or are slightly undervalued. Additionally to this, the powder density of the powder layer in the simulation is less. Both facts agree to the approach that the powder density of metal powder is reducing at the layer coating by a wiper.densityoftherealpowderis measuredaccordingtoFurthermore, the ratio of the thermal conductivity of the loose simulated powder to a solid body out of the same material ascertained in the plane of the layer, which is the plane crucial for the process, is given asThe ratio is between 13 % and 42 % (column 4 in table 7). A look to the ratio of the thermal conductivities between powder and solid body in the simulation near room temperature shows too high values since the thermal conductivity of powder should be around ten times of the atmosphere gas, arising ~0.2 W/(m·K) for argon [Rombouts et al., 2005].

If we increase the distance between the spheres, the thermal conductivity would be reduced, but also the powder density. For example, a distance of 5 µm in the powder bed Al_20+40_4-6 leads to a ratio of thermal conductivities of 6 %, however also to a powder density of around 40%, which is much smaller than the density of 54 % in the real loose powder. Thus, the thermal conductivity in our simulation does not reach the values in the literature [Zeng et al., 2012]. As one reason we assume the numerical discretization between the particles, which result in overestimated particle connections. Additionally, other aspects like oxide layers, which cannot yet be displayed in our simulation model, can lead to a reduced thermal conductivity in the powder material.The properties of the powder bed in the process chamber of the LBM machines can hardly be determined experimentally. On the one hand the verification methods like camera imaging have not the resolution to resolve single particles.

On the one hand the verification methods like camera imaging have not the resolution to resolve single particles. On the other hand methods like scanning electron microscope or laser scanning microscope cannot be set up in the process chamber. Regarding the currently available verification methods for the powder bed in commercial machines like the measurement of the powder density, we state a good agreement of the real and the modeled properties of the powder beds. The powder density of the powder mixtures, which is the most important property, accords with the real loose powder and the powder layer is satisfactorily represented. Unfortunately, the thermal conductivity of the powder layer is too high at room temperature. But it highly increases with the temperature [Zehner, 1972], especially by effects of sintering.

Unfortunately, the thermal conductivity of the powder layer is too high at room temperature. But it highly increases with the temperature [Zehner, 1972], especially by effects of sintering. Besides, the capillary actions between the particles and the melt [Gürtler et al., 2013] drastically change the properties of the powder bed and consequently the thermal conductivity at the main focus of the process. Thus, the additional distance between the particles can be considered as correctly chosen.Process SetupFor the investigation of the influence of different powder beds on the process a workspace of 400 x 400 x 200 µm is analyzed. The temperature in this box is set to 500 K at the start time. The material and energy can flow through the boundaries, except for the bottom side. The temperature under an 80 µm thick metal plate is set to constantly 500 K, since that allows displaying the high energy loss by the already molten and solidified layers.

The temperature under an 80 µm thick metal plate is set to constantly 500 K, since that allows displaying the high energy loss by the already molten and solidified layers. The length of each line processed with the laser beam in the powder bed is 300 µm long. The first line is 50 µm beside the middle of the workspace, the second is in the middle and the last line is located 50 µm further. As in the experiments, the line spacing is 50 µm, the feed rate is 0.75 m/s, the laser power is 100 W, the wavelength is 1064 nm and the beam waist is 10 µm (table 2). The simulations are done for the presented regular powder beds and for imperfect powder beds, in which the large particles along the diagonal of the workspace are missing, see figure 5. Thus we can investigate the influence of defects in the powder bed on the process.a) Al_20+32-40d) Al_20+40_3-7b) Al_20+40-50e) Al_20+40_4-6c) Al_20+50-63f) Al_20+40_5-5Fig. 5.

Thus we can investigate the influence of defects in the powder bed on the process.a) Al_20+32-40d) Al_20+40_3-7b) Al_20+40-50e) Al_20+40_4-6c) Al_20+50-63f) Al_20+40_5-5Fig. 5. Illustration of the different powder beds with missing large particles along thediagonal of the workspace (a-f)Results and DiscussionExperimental resultsFor comparison experiments are carried out with the same parameters as in the simulations and a layer thickness of 30 µm. Cubes with 5 mm edge length are built on the machine SLM 50 by ReaLizer GmbH and their relative densities are investigated, see figure 6. The values of these results are split in three different groups. For later comparison to the simulation we define a green quality class for values over 98 %, yellow for 97-98 % and red for values lower than 97 % relative density. Accordingly, the green group represents more dense work pieces than the orange ones and red for worse subjects.

7. Illustration of three molten scan lines for the powder beds Al_20+40_4-6 withmissing large particles along the diagonal of the workspaceIn the next step we investigate the thickness of the resulting layers. An outcome can be taken from figure 7a where the three molten tracks of the inhomogeneous powder beds Al_20+40-50 and Al_20+40_4-6 are shown. A groove is formed along the missing spheres, which is distinctively visible in figure 7b.

A groove is formed along the missing spheres, which is distinctively visible in figure 7b. In the following we analyze this by various measurements, starting with a theoretical analysis.Table 8: Theoretical layer thickness for regular and inhomogeneous powder beds calculated usingthe powder density and the relative height of the layer at defectsPowder nameAl_20+32-40 Al_20+40-50 Al_20+50-63 Al_20+40_3-7 Al_20+40_4-6 Al_20+40_5-5Theoretical height in µm 17.2 23.4 26.1 21.9 25.6 21.7Height at missing spheres in µm 9.5 2.4 2.6 14.2 10.2 10.8Relative height0.55 0.10 0.10 0.65 0.40 0.50Using the powder height and powder density (from table 6), we calculate the theoretical thickness of the powder layers in the unit cells of the particles. In the second column of table 8 the theoretical heights of the regular layer are shown for the different powder beds.

In the second column of table 8 the theoretical heights of the regular layer are shown for the different powder beds. In the next column the theoretical heights of the powder beds, in which the large particles at the diagonal are missing, are determined in the calculations by removing the volume of the missing sphere in the unit cells.

In the last column the relative height , which is the ratio between the inhomogeneous and regular powder bed, is calculated to give a better comparison:with the area at the bottom of the unit cell and previously defined variables and .Table 9: Simulated layer thickness after solidification in the middle of the workspace for theregular and inhomogeneous powder beds and the relative height of the layer at defectsPowder nameAl_20+32-40 Al_20+40-50 Al_20+50-63 Al_20+40_3-7 Al_20+40_4-6 Al_20+40_5-5Simulated height in µm 12 24 26 22 22 18Height at missing spheres in µm 4 2 0 0 0 5Relative height0.33 0.08 0.00 0.00 0.00 0.28The same analysis is carried out for the heights and within the simulations (table 9). We measure the layer thickness in the middle unit cell of the workspace, which means in the middle of the second scan line.

We measure the layer thickness in the middle unit cell of the workspace, which means in the middle of the second scan line. The measured minimum height inside the unit cell of the particles is used as the layer thickness to avoid influences of the geometry of the particle arrangement. The simulated layer heights are similar or smaller than those calculated theoretically. The heights in the simulations with missing spheres vary between 0 and 5 µm. This means that the relative heights , which are defined as the ratio between the simulated heights of the regular powder beds to those of the inhomogeneous powder beds at the position of the imperfection, reach 28 % maximum:Both height measurements for powders with large particles are zero.

As a first explanation for this phenomenon we infer that powder beds with small particles can compensate defects in the powder bed better than powder beds with large particles.For the following analysis we calculate the ratio of the simulated to the theoreticallayer heights of the inhomogeneous powder beds (column 3 in table 8 and 9):The values are between 0 and 85 %, depicted by bright colors in figure 8. This measurement shows the compensation of defects in the unit cells by melt pool dynamics. As in the experiments (compare figure 6), we can see three quality groups, the red quality class at zero, the bright orange group at around half the value of the green class, which is at 85 %. Thereby we can easily distinguish the powder mixtures for the suitability in LBM.

Thereby we can easily distinguish the powder mixtures for the suitability in LBM. The green group stands for a better compensation of defects in the powder bed and consequently for a better resulting work piece than the orange one and red means a worse quality.o i t a R0,90 0,80 0,70 0,60 0,50 0,40 0,30 0,20 0,10 0,002 4 , 01 6 , 05 8 , 03 8 , 00 0 , 00 0 , 00 0 , 00 0 , 00 0 , 00 0 , 08 2 , 06 5 , 0Ratio of sim. to theo. heights at defectsRatio of relative heightsFig. 8. Ratio of layer heights between simulation and theory at missing spheres in the powder bed (bright colors) and ratio of relative heights between simulation and theory (dark colors): They are split in three quality classes (green: >75 %; orange: 25-75 %; red: <25 %), similar to the experimental results (The red values are zero.).If we compare the relative layer heights of the simulations to those of the theoretical analysis, we obtain similar results.

Therefore, the ratio of the simulated relative height to the theoretical one is calculated (column 4 in table 8 and 9):Subsequently we get a completely normalized criterion to evaluate the quality of the powder mixtures for LBM. The results are displayed by dark colors in figure 8. We split them in three quality classes with similar values as for the ratio of layer heights. The green group signifies a better conformance of the theoretical defects compensation and again a better process result than the orange one and red agrees worse.For both ratios the powder mixtures are separated into the same quality classes. The powder Al_20+50-63, Al_20+40_3-7 and Al_20+40_4-6 have got the worst results and are in the red class. The mixtures Al_20+32-40 and Al_20+40_5-5 have got ratios of around 0.5 and are in the orange class. The green quality class consists of the powder mixture Al_20+40-50 and achieves a value of over 80 %.

The green quality class consists of the powder mixture Al_20+40-50 and achieves a value of over 80 %. This quality order agrees completely with the findings of the experiments, see figure 6.Through our simulative fluid-dynamic analysis we can confirm the assumption that powder beds with more and smaller particles compensate defects in the powder bed better by means of melt pool dynamics. Thus, the surface is smoother and the following layer has a more even thickness. As a consequence, the process is producing more actually and there are fewer defects in the work piece, which is confirmed by the experiments.Analysis of the simulated melt poolIn this section the melt pool shape is analysed for a deeper process understanding.

Itsvolume is measured at the time when the laser is located in the middle of the second track.Table 10: Melt pool volume in the middle of the process for regular powder beds and powderbeds with defects and their relative volumePowder nameAl_20+32-40 Al_20+40-50 Al_20+50-63 Al_20+40_3-7 Al_20+40_4-6 Al_20+40_5-5Volume at reg. beds in µm³ 30.4·103 33.2·103 73.5·103 34.8·103 42.0·103 56.7·103Volume at defects in µm³ 11.8·103 7.0·103 37.2·103 9.6·103 18.0·103 30.4·103Relative volume 0.39 0.21 0.51 0.28 0.48 0.54The melt pool volume is reduced by defects in the powder layer to a ratio of 21-54 % (table 10). The decrease occurs since the laser beam energy is directly conducted by the solid layers (metal plate) at the bottom because of the missing heat accumulation of the completepowder layer.

The decrease occurs since the laser beam energy is directly conducted by the solid layers (metal plate) at the bottom because of the missing heat accumulation of the completepowder layer. However, the outcomes show no clear trends to discriminate the powder beds for process stability.Table 11: Melt pool depth in the middle of the laser spot in the middle of the process for regularpowder beds and powder beds with defects and their relative depthPowder nameAl_20+32-40 Al_20+40-50 Al_20+50-63 Al_20+40_3-7 Al_20+40_4-6 Al_20+40_5-5Depth of reg. beds in µm 23.5 22.4 18.0 21.6 29.7 27.9Depth at defects in µm 12.0 13.3 7.2 13.3 22.8 23.3Relative depth0.51 0.59 0.40 0.62 0.79 0.84Because in LBM there are no effects similar to the keyhole at laser beam welding, especially at these process parameters, the melt pool depths can be measured in the middle of the laser spot when the laser beam is in the middle of the second line (table 11).

The ratios of the melt pool depths of the inhomogeneous to the regular powder beds vary within a range of 40-84 %. These results show a tendency to better preservation of the melt pool depth against defects by more and smaller particles. Hence, this outcome agrees with the evidence of the consolidated layer heights shown in the section before.Summary and OutlookA numerical method at the powder scale is used in this paper to investigate the LBM process. We investigate the influence of various particle size mixtures of the material AlSi10Mg and defects in the powder bed onto the melt pool and the porosity formation.Therefore, we verify the powder representation of our simulation model and ensure that the used powder configurations agree with reality.

Despite the thermal conductivities of our powders at room temperature are overestimated due to geometrical constraints, we point out that their thermal conductivity at much higher process temperatures can safely be assumed to match the experimental values more closely. The powder densities of our powders agree well with the experimental values.First we examine the resulting layer thickness. The measured heights in the simulations normalized to the theoretical calculated heights of the compacted powder layer show a good correlation to the particle size distributions. Furthermore, we investigate the melt pool volume and height. We determine that the powder mixtures with more and smaller particles compensate defects in powder beds better. This leads to a more stable LBM process, which is in accordance with the results of our experimental investigations.

This leads to a more stable LBM process, which is in accordance with the results of our experimental investigations. In sum, the findings of the current study clearly demonstrate the importance of error free layer coating for porosity formation and overall quality of the LBM process.Further research will focus on the propagation of defects through several layers. This willenable us to analyze the resulting powder bed whether a compensation of defects is possible.AcknowledgmentsThe authors gratefully acknowledge funding of the Erlangen Graduate School in Advanced Optical Technologies (SAOT) by the German Research Foundation (DFG) in the framework of the German excellence initiative and funding of the German Federal Ministry of Education and Research by the funding program Collaborative Research Center SFB814 (Additive Manufacturing), project A5. Work is performed with support of the Ministry of Education and Science of the Russian Federation, the state contract No.

lid Freeform Fabrication 2019: Proceedings of the 30th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperINFLUENCE OF POWDER PARTICLE SIZE DISTRIBUTION ON THE PRINTABILITY OF PURE COPPER FOR SELECTIVE LASER MELTINGM.(cid:3)Sinico1,2, G. Cogo2, M. Benettoni2, I. Calliari3, and A. Pepato21 Department of Mechanical Engineering, KU Leuven, 3001 Leuven, BE 2 National Institute for Nuclear Physics - Padova Division, 35131 Padova, IT 3 Department of Industrial Engineering, University of Padova, 35131 Padova, ITAbstractThis work investigates the use of fine Cu powder, with ~ 20 vol% smaller than 15 μm size, for the selective laser melting process. Cubes reaching > 98 % density are produced at relative low laser output (175 W) and characterized. After the selection of a proper combination of laser scan parameters, the properties of fabricated parts are briefly studied through profilometry and tensile tests.

After the selection of a proper combination of laser scan parameters, the properties of fabricated parts are briefly studied through profilometry and tensile tests. Finally, a voluminous demo component for high-energy physics is manufactured to stress-test the employed SLM machine. Even though unmolten particles and lack of fusion defects are still present in the produced specimens, the investigated approach confirms that powder selection can have a huge influence on the processability of materials with high reflectivity towards near-infrared irradiation.IntroductionSelective Laser Melting, also called Laser Powder Bed Fusion, is an Additive Manufacturing technique designed to use a high power laser to melt and fuse metallic powders together in a layer-by-layer fashion.

The process has seen a considerable industrial adoption during recent years, thanks to the possibility to produce freeform near-net shape components with low material waste and reduced design-to-production time [1]. Despite the advantages of this fabrication technique, limitations are still present with one main restriction being the available material palette. Only few materials are currently used at an industrial scale, with many more studied at academic level, due to the fact that SLM is a complex multiphysics process, and that the successful SLM fabrication is dependent on a considerable amount of variables. Those variables, with a sub-section of them being the SLM process parameters, need to be fine-tuned to obtain dense and accurate parts with suitable mechanical properties, but the identification of this optimal process windows is often time and resource intensive. Moreover, the material needs to be available in powder form with a specific granulometry and high sphericity.

Moreover, the material needs to be available in powder form with a specific granulometry and high sphericity. One of the materials of interest that still poses some obstacles for its effective SLM fabrication is pure copper. Copper not only suffers from its most known properties, with a higher tendency of melt pool instability during SLM due to the rapid dissipation of the deposited heat input. It is also one of the materials with the highest reflectivity towards the near-infrared radiation, which corresponds to the emitted wavelength region of commonly adopted lasers in SLM systems (continuous wave fiber lasers). This poses serious troubles on its effective SLM fabrication, since only a fraction of the irradiated power is absorbed and contributes to the melting process, leading to a porous component with many unmolten particles.

Various researchers during the years have suggested diversified approaches to successfully fabricate this difficult material, and the most common ones to date are to use copper alloys or a mixture of powders with copper particles and other elements to increase the absorbance in the infrared at the cost of electrical and thermal conductivity [2]. Other solutions have been the implementation of green or blue lasers in SLM machines [3], since the reflectivity of copper decreases with wavelength from the infrared through ultraviolet regions, or the use of powerful fiber lasers up to 2000 W [4,5,6]. None of the latter proposed solutions have been industrially accepted, on account of the fact that the common available SLM machines do not support these kind of options and should be modified individually at a cost.

All the attempts at a lower laser power [7,8] using CO2 or standard fiber lasers have failed, never reaching a density of the produced pure copper parts higher than 90 %.657The SLM of pure copper material could unlock the possibility to produce innovative heat exchangers with enhanced thermal conductivity [1], complex induction coils with increased electrical conductivity [9], and many other components like heat sinks in injection molding, copper anodes for fuel cells, and miscellaneous parts [10] where thermal and electrical properties are more of a concern. This demand arises not only from the industry, but also from different fields of academic research like experimental physics or aerospace, where the adoption of SLM could lower the design-to-production time of unique components that are nowadays produced with several steps of standard manufacturing processes, will all the connected limitations.

The AM of copper in these research fields has already being spotted as a solution e.g. for components in the DEMO (DEMOnstration power plant) and ITER (International Thermonuclear Experimental Reactor) fusion power plants [11], or 3D printed rocket thrust chambers [10].In this work, the influence of powder particle size distribution is studied and correlated to the SLM processability of pure copper. It is well known [12,13] that powder granulometry influences the SLM printability: fine spherical particles with a wide enough size distribution are reported to be easier to be molten in respect to coarser ones, generally increasing powder bed packing density and favoring part densification. In addition, the particle size should be tuned on the employed SLM layer thickness to promote an homogeneous spread of the powder layer.

In addition, the particle size should be tuned on the employed SLM layer thickness to promote an homogeneous spread of the powder layer. On the other hand, fine powders also present lower flowability, caused by their higher tendency towards cohesion and agglomeration due to stronger inter-particle attractive forces, and an increased chemical reactivity, e.g. higher propensity to oxidation. Consequently, there should always be a balance between the positive (e.g. easier melting process) and negative (e.g. decreased flowability) effects of introducing finer particles in the SLM process, to maximize the quality of the printed part. As depicted before, machine modifications (green-blue lasers, high power fiber lasers) and raw material modifications (alloying elements) have been already explored as possible solutions for copper SLM production, while the combination of fine powder with a thin layer thickness has never been reported in literature.

Our hypothesis is that, despite the high reflectivity, this combination of favorable conditions could ease the melting process of pure copper, allowing to decrease the total amount of porosities in the final printed part. To test our hypothesis, one standard pure copper powder for SLM and one engineered distribution of fine copper powder have been acquired and processed, with the results hereby reported in the following essay.Powder characterizationTwo gas atomized pure copper powders (99.8 wt% Cu) with different powder particle distributions have been purchased from LPW Technology Ltd. The first distribution, hereby called Cu 15-45, is the standard size distribution being sold by the supplier. The second distribution, hereby called Cu 10-35, has been engineered in collaboration with the supplier to meet defined constraints. Specifically, the engineered powder should be suitable to be deposited in building layers of 20 μm.

Specifically, the engineered powder should be suitable to be deposited in building layers of 20 μm. Therefore, the D50 of the powder distribution is requested to be around the employed layer thickness, while the D90 should be below the effective layer thickness. D50 and D90 are the values of the particle diameter at 50 % and 90 % in the cumulative distribution, respectively. The effective layer thickness teff is calculated assuming a powder layer density of 60 % [13], and corresponds to 33.3 μm. Laser diffraction analyses have been provided by the supplier after the atomization of the two powders (Figure 1), confirming that the Cu 10-35 engineered distribution meets the requirements. Moreover, the presence of a tail of smaller particles, with a volume below 15 μm of 21.31 %, is considered positive and will generally help to fill the voids between coarser particles and consequently increase the powder bed packing density and ease the melting process [12].

Scanning Electron Microscopy (SEM) inspection has also been performed, and is reported in Figure 1. From the captured SEM images it can be noticed that the majority of particles exhibits a spherical shape, along with a limited amount of irregularly shaped particles. The standard distribution Cu 15-45 seems to present a lower mean sphericity and a slightly discernible higher amount of satellites compared the engineered distribution Cu 10-35.to65815Volume Below (15) µm 4.60 %Dv {10) 17.0 µm10015Volume Below (1S) µm 21.

The baseline correction for the measurements has been performed on the provided barium sulfate white board. Results for the two copper distributions (Figure 2) are compared with results acquired on stainless steel 316L and bronze CuSn10 powders gently provided by the company SISMA S.p.A.80Nd:YAG emission wavelength ~107075706560~ 055Q) uC .mu Q) ~ 0:::50454035!~35.3%30252015- SS 3 16 L 15-45 µm - - Bron ze 10-35 µm - - cu 99 .8 10-35 µm - - c u 99 .8 15-45 µm10400500600700 800 900 1000 1100 1200 1300 1400Wave length , A [nm]Figure 2. Laser absorption measurements via spectrophotometer of the selected pure copper powders, and of stainless steel 316L and bronze CuSn10 powders for comparison.The optical reflectance of the used copper powder bed is ~70 % at a wavelength of 1070 nm (near-infrared) for both Cu distributions.

This result is comparable to previous reported laser absorption measurements [4,14] and confirms one of the main challenges of SLM copper production. It is worth noticing how the engineered Cu 10-35 powder actually presents a slighter higher reflectance, probably due to a higher powder bed packing 65910050010,000.0~Q)E ::, gQ) > -~ :i E ::, (.)density which could lower the multiple internal reflections and subsequent absorption of the optical light between separate particles. The difference in reflectance is better compared with the bronze CuSn10 powder, where the presence of Sn in 10 wt % in the alloy is able to decrease the reflectance from ~70 % to ~53 %, further evincing why copper alloys have been more successfully processed with the current market of SLM machines [2,15,16].SLM setup and experimental methodologyThe performance of the two copper powders have been evaluated with the production of specimens through a SISMA mysint100 PM SLM machine.

5bar @ 35 UminInert gas consumption<0,3 Umin @ 0,5% 0202 conce ntration02, <lOOppmMachine dimensions1390 mm x 777 mm x 1600 mm ILxWxHINet weight650 kgCubic specimens 10x10x10 mm (example in Figure 3a) have been produced with both Cu 15-45 and Cu 10-35 powders, at a fixed laser power of 175 W. The laser power has been fixed to that value to maximize the surface energy density Es; the maximum power of 200 W has not been used to prevent precocious wear of the laser source, given the high scheduled amount of experiments.a)b)Figure 3. a) Example of 27 10x10x10 mm specimens arrangement on the build platform ((cid:4231)platform = 99 mm) for SLM parameters optimization study; b) example of horizontal tensile test specimens arrangement.660Layer thickness has been as well fixed to a value of 20 μm, whereas laser scanning speed, hatch spacing and scanning pattern have been varied.

No powder flowability issues have been encountered while using the Cu 10-35 fine powder distribution, and a standard rubber recoating system has been enough to provide homogeneous powder layers during SLM processing in nitrogen atmosphere. Relative density measurements have been performed, after careful cutting and polishing of the cubes, with optical microscopy and, for selected specimens, SEM inspection. Moreover, chemical etching through a mixture of 2.5 g FeCl3 – 1 ml HCl – 100 ml EtOH have been employed to highlight observed lack of fusion porosities. All the samples have been built on top of a specifically designed copper OFHC platform. After SLM parameters optimization, specimens for tensile tests (example in Figure 3b) and surface roughness measurements have been fabricated with the Cu 10-35 powder batch.

After SLM parameters optimization, specimens for tensile tests (example in Figure 3b) and surface roughness measurements have been fabricated with the Cu 10-35 powder batch. The tensile samples have been designed following the ASTM E8M - 16a standard (test piece diameter do = 4 mm) and tested in as-build condition on a MTS 858 Mini Bionix II with a displacement rate of 2 mm/min. A MTS extensometer 632.29F-30 with a gage length of 5 mm was used to measure the strain during the tensile tests.

A MTS extensometer 632.29F-30 with a gage length of 5 mm was used to measure the strain during the tensile tests. Finally, an optical profiler Sensofar PLu neon has been employed for the roughness analyses, following the ISO 4287-1997 standard.Results and discussionSLM processingThe SLM parameters optimization is summarized in Figure 4, where obtained relative densities of the cubic specimens after optical measurement are compared between the two Cu powders and plotted as a function of volumetric energy density ((cid:1831)(cid:3049) = (cid:1842) ((cid:1874) × (cid:1860) × (cid:1872)) with P laser power, v laser scanning speed, h hatch spacing and t layer thickness). A full analysis of the dependence of relative density on the different set of parameters, e.g. v, h, scanning pattern, is scheduled to be published in a full journal paper.(cid:932)10098■■SEM surface Cu 15-45a: ~ (/) C Q) 1J Q) > ~ ai [J:'.--■~--·9694 I__ _ ___..

On the other hand, the selected Cu powder presents an impressive improvement in densification, with a maximum value of 98.1 %. From SEM inspection of top surfaces, the Cu 15-45 specimens can be described as a porous structure with melting beads and almost non-distinguishable laser line tracks. Cu 10-35 specimens, instead, have a stable fusion and more continuous line tracks, even if some balling, recesses and weld interruptions remain. To identify the causes of the remaining porosity content in the Cu 10-35 samples, etching has been performed and an example can be seen in Figure 5.661Lack of fusionFigure 5. Optical inspection of a selected Cu 10-35 specimen (from Figure 3) after etching of a Z cross section, i.e. building direction; lack of fusion drawing courtesy of [17].It’s clear from the etched sample that lack of fusion still compromises the successful SLM production of pure copper.

building direction; lack of fusion drawing courtesy of [17].It’s clear from the etched sample that lack of fusion still compromises the successful SLM production of pure copper. Even if a higher relative porosity is achieved with the selected powder, semi-molten particles and “knit” weld lines between subsequent layers are not completely avoided. Lower mechanical properties, e.g. strength and ductility, in the Z direction (anisotropy) and an overall low fatigue life of components are to be expected. Tests are ongoing to evaluate the electrical and thermal conductivity of the samples with the highest relative density in respect to the IACS standard. We believe that parts produced with the optimized set of parameters can’t be certified for applications that might require intensive mechanical load, but could nevertheless possibly be adopted for specific applications where electrical and thermal conductivity are more of a concern.

It remains anyhow impossible at this stage to draw conclusions before the complete characterization of the samples. As a final remark, it is imperative that a more in-depth analysis of the total porosity content should be done, with the comparison of the obtain values from optical inspection against Archimedes method and possibly micro Computed Tomography. Although the Struers guidelines [18] have been followed during the metallographic preparation of the specimens, the possibility of deformation caused by the grinding and polishing steps could partially hide the smaller pores and thus falsely lower the porosity content. That said, the investigated approach confirms nonetheless that powder selection can have a huge influence on the SLM processability of pure copper, and powder morphological properties like particle size distribution should preferably always be tuned beforehand on the SLM system and process parameters window, e.g.

layer thickness, that are going to be employed.Mechanical tests and surface roughness measurementsThe mechanical tests, performed on series of 8 samples built both on the XY plane (Figure 3b) and on theZ direction (building direction), are summarized in Table 2.Table 2.

Results of the mechanical tests on the SLM fabricate Cu samples, reference values of Cu C11000 properties from [19] for comparison.Tensile strength [MPa]0.2% yield strength [MPa]Modulus of elasticity [GPa]Elongation at break [%]SLM Copper, XY direction207 ± 26157 ± 23133 ± 219.83 ± 0.83SLM Copper, Z direction Reference, C11000 electrolytic tough-pitch Cu, cold-worked (H04 temper) Reference, C11000 electrolytic tough-pitch Cu, hot-rolled130 ± 23345220125 ± 193106983 ± 16112 - 130112 - 1302.48 ± 0.431245Note: all data points with their corresponding standard deviation.662The surface of the specimens was not treated after SLM fabrication (as-built), and only the supports were detached from the parts. The parameter set used for these samples provides a corresponding Ev of ~583 J/mm3, with an expected relative density of 97.1 % (Figure 4).

The parameter set used for these samples provides a corresponding Ev of ~583 J/mm3, with an expected relative density of 97.1 % (Figure 4). It was decided to avoid the parameter combinations at higher Ev since their volumetric material fabrication rate was too low to be applicable at an industrial scale. As per our knowledge, the only reported mechanical test on pure copper fabricated by SLM was 149 MPa of ultimate tensile strength for a relative density of the sample of 88.1 % (XY direction) [7]. In this work, a tensile strength of 207 MPa is achieved for the horizontal XY direction, further corroborating the improvement of our proposed approach. Overall, the obtained XY properties are analogous or superior than C11000 electrolytic tough-pitch Cu in hot-rolled condition. The major drawback is solely the percentage of elongation, which -mostly due to the remaining porosities and lack of fusion defects- halts at a value of ~10 %.

The major drawback is solely the percentage of elongation, which -mostly due to the remaining porosities and lack of fusion defects- halts at a value of ~10 %. This behavior is more accentuated for the samples built in the Z direction, where all measured mechanical attributes deteriorate considerably. More specifically, the tensile strength drops to a value of 130 MPa while the percentage of elongation decreases to ~2.5 %. As predicted, lack of fusion is a typical defect that diminishes the bonding between SLM layers, and therefore lowers the mechanical properties of printed parts in the Z direction. Fabricated objects will experience extensive anisotropy and an expected low fatigue life. It must be specified that the influence of minor alloying elements (0.2 wt%), crystal structure, crystal orientation, phases or strain in the material are not accounted for and they could as well have a considerable impact on the reported measurements.

It is well known [4,8] that SLM copper commonly presents a fine grained microstructure and might display a crystallographic texture due to the generated temperature gradients within the melt pools. Both of these microstructural contributions should be further investigated to better understand and predict the mechanical behavior of SLM copper material.Finally, the surface roughness analyses performed on a rhombohedral specimen are shared on Figure 6. The specimen is produced with the same parameters used for the mechanical test samples. Each reported roughness measurement has been repeated at different locations on each surface, taking care that the profiles are measured perpendicular to the typical lay lines featured on SLM surfaces, i.e. the laser line tracks.20 ~ - - - - - - - - - - - - - - - - - -~ 30019E 18; 17 Ct'. ::l 16 Q) ~ 15 Cl :, e 14 ~ ~ 13 :, (I) 12-■- - Ra value - _,,_ - R2 value, •---------r, f .•····•·1·· ..... 1····I280 260 E 2.

::l 16 Q) ~ 15 Cl :, e 14 ~ ~ 13 :, (I) 12-■- - Ra value - _,,_ - R2 value, •---------r, f .•····•·1·· ..... 1····I280 260 E 2. t2_240220::lQ)200 ~ Cl :, 180 eQ)160 ~ :, 140 C/J90°Top surface 0°135°1112010 ~ -~ - - - -~ - - - -~ - - - -~ -~ 10004590135SL M fabrication angle [0Figure 6. Surface roughness measurements of the Cu 10-35 rhomboid at different SLM fabrication angle, all data points reported with their standard deviation.The Ra results, with 12.21 ± 0.25 μm for the top surface, are comparable with the standard surface quality achieved by common SLM machines [20] and therefore reflect the improved melting and higher relative density of the Cu 10-35 samples. The trend of the surface roughness at different angles is also reasonable, with a decrease in surface quality for sloped surfaces where the staircase effect comes into play.

The trend of the surface roughness at different angles is also reasonable, with a decrease in surface quality for sloped surfaces where the staircase effect comes into play. In addition, for the 135° surface, the combination of the staircase effect with being a downfacing region considerably worsens the surface condition. However, more information can be retrieved if attention is put on the Rz parameter and not on the arithmetical mean deviation parameter Ra, that is by definition an average. It is then noticed how actually Rz, or the maximum height of the roughness profile, is mildly higher than regular [20] probably due to the fact that -even with an 663improved melting- some balling, surface porosities and laser tracks irregularities remain in the samples.

Those features, as shown before in the SEM image (Figure 4), will determine few larger peaks and valleys in the roughness profile which consequently reflects better on the maximum height parameter.SLM stress-test: building a demo component for high-energy physics applicationGiven the high reflectivity of pure copper (Figure 2) the risk of laser back-reflection is considerably high and has been recently reported in literature [4]. Back reflections can create new focal points inside the optical system during processing and thus damage the optical mirrors, compromising the SLM system. Although this issue is more prone to occur at higher laser powers than the one used in this work (175 W), it is withal deemed paramount to check if the SLM machine can withstand production of massive copper parts.

A demo component for high-energy physics is consequently selected for production as the final test-sample of our research; the detailed description and function of the part are not the main subject of this paper and, hence, they will be just briefly described. The part is a beam dump designed to stop the proton beam of a cyclotron at the end of a beamline [21]. The beam dump has to guarantee the full power heat removal in a safe way fulfilling all the design specifications as well as radio protection limits (examples of beam dump designs in [22,23]). The object can be conceived as a heat exchanger, where the protons are initially decelerated by bulk copper material, and this material must be cooled down to avoid the part to reach its melting temperature and/or have material vaporization inside the beamline.

In details, for the proposed application, the component must be able to dissipate ~30 kW of a Gaussian 70 MeV beam power deposited in a beam spot size of 21.4 cm2 surface area (RMS, ±3 (cid:305)). That is equivalent to 1.4 kW/cm2 of energy that should be removed by a suitable cooling system. The following sketch helps to recap the main concepts./H, O or He-gas1S0-K 100 ;\.I flangeVacuuma)b)30 kW Proton Beam on 21.4 cmA2Figure 7. Explanatory sketch of the beam dump purpose, a) beamline end with the component in orange plus b) details of the design constraints.To perform the task, the beam dump should have a front portion, facing the beam, composed of bulk dense copper to effectively stop the incoming protons, and a second back portion where the cooling system acts to cool down the front.

The thickness of the front portion, dependent on the stopping power of the material [24] and on the incident beam properties, is estimated to be ~8 mm for Cu OFHC (Monte Carlo SRIM-2013 simulation [25]). Since, as previously reported, the SLM of pure copper still cannot output fully dense parts, the beam dump front side is directly obtained from the Cu OFHC platform, where the higher purity and density of the material enhance its thermal conductivity and stopping power. Only the cooling system, designed in a spiral way to remove the generated heat from the center of the component all the way to the outer sides, has been 3D printed on top of the Cu OFHC platform. Figure 8a portraits the main idea, with 10 mm (8*1.25 safety factor) of bulk high purity copper in the front portion. The AM design freedom is adopted for the conformal cooling spiral channel, as well as -potentially- the implementation of a metallic foam e.g.

The AM design freedom is adopted for the conformal cooling spiral channel, as well as -potentially- the implementation of a metallic foam e.g. metallic struts inside the channel itself to increase the turbulence of the cooling liquid. This design and manufacturing approach saves production steps and, if the664enhanced thermal efficiency is proved, could be a viable way to create a smaller and lighter beam dump compared to the conventionally manufactured ones presented in scientific literature [22,23]. The part has been fabricated with a production set of SLM parameters, to keep the overall production time at a reasonable level. The corresponding Ev of ~365 J/mm3 is consequently lower than the obtained one for the tensile test specimens. The 3D printed cooling system has a total volume of 88392 mm3, with a printing height Z of 32 mm.

The 3D printed cooling system has a total volume of 88392 mm3, with a printing height Z of 32 mm. The top surface of the OFHC platform has been sandblasted prior to production to avoid laser back reflections from the polished surface of the platform itself, given the 90.1 % reflectance for polished or electroplated pure copper surfaces [19].a)b)SLM 3D printed Cu channeOFHC from Printing PlatforFigure 8. a) Side view of the CAD design, highlighting the 10 mm bulk Cu OFHC from the building platform; b) beam dump after successful SLM production, with its spiral cooling system.In Figure 8b, a photo of the beam dump after SLM manufacturing is shown. The part was successfully produced in one single job of ~75 hours. No major drawbacks were noticed during the SLM phase, and after the positive result of this stress-test the same SISMA mysint100 PM machine was later employed to produce other Cu showcases and demo components.

We want to disclose that the beam dump case study hereby described has been designed as a demo part as well, based on real data from the SPES cyclotron project [21]. The goal was to highlight AM and specifically SLM copper production potential for high-energy physics applications, and further design iterations, simulations and tests will be necessary to develop the final product.ConclusionsPure copper specimens and demo-components have been fabricated by Selective Laser Melting reaching a relative density of 98.1 %. A novel approach has been suggested to overcome the difficulties encountered in the AM laser production of copper, combining the use of a fine Cu powder (in the range of 10-35 μm) and a small layer thickness to ease the melting process with standard fiber lasers and reduce the lack of fusion defects.

Comparing the obtained results against standard size powder (15-45 μm) processed at the same combinations of SLM parameters, it has been demonstrated how powder selection is the critical factor and should always be tuned beforehand based on the SLM machine, processing conditions, and material. The proposed approach yet cannot achieve full density parts, with lack of fusion defects still present in the processed samples. Anyhow, the diminished total porosity content confirms that the use of powders with a PSD skewed towards finer particles is effective to manufacture relatively high density pure copper parts at a laser power lower than 200 W. Moreover, complete mechanical properties have been reported for SLM fabricated pure copper for the first time in both horizontal and building direction. The measured properties, combined with the good obtained as-build surface quality, confirm once more the initial hypothesis of this work.

The measured properties, combined with the good obtained as-build surface quality, confirm once more the initial hypothesis of this work. Analogous or superior mechanical strength than C11000 electrolytic tough-pitch Cu in hot-rolled condition can be achieved for the horizontal fabricated specimens, while in the building direction the tensile strength and elongation at break drop due to the lack of fusion between building layers. In view of the anisotropic behavior, the produced parts probably cannot be certified for applications that might require intensive mechanical load, but could nevertheless possibly be adopted for specific functions where electrical and thermal conductivity are more of a concern.665Hence, a voluminous demo component for high-energy physics has been fabricated as a proof-of-concept of a possible thermal application and to stress-test the SLM machine in use.

The part was successfully built in ~75 hours, and no critical laser back-reflection or any complication occurred during the build job. Further studies are ongoing to obtain a more in-depth analysis of the total porosity content in correlation to the applied scanning parameters and strategies; furthermore, the influence of minor alloying elements, crystal structure, crystal orientation, phases or strain in the material should be investigated to better understand their contribution to the measured mechanical properties. Finally, the electrical and thermal conductivity of the fabricated copper samples with the highest relative density are being measured, and will be reported and discussed in a forthcoming journal paper.Acknowledgementsthe project 5687-1-2121-2015 POR FSE 2014-2020 Veneto “Implementation and characterization of advanced additive manufacturing techniques, applied to the production cycle of the mechanical industry in the Veneto Region”.

The industrial partners SISMA S.p.A. and Officina Dei Materiali SaS are kindly acknowledged for the continuous guidance and training throughout the project and for giving access to their laboratories and equipment. We would also like to express our gratitude to Eng. Daniele Rigon and Prof. Giovanni Meneghetti, who performed the Department of Industrial Engineering, University of Padova. Moreover, assistance provided by Eng. Matteo Bellin for the roughness measurement was greatly appreciated. Finally, special thanks to Dr. Juan Esposito for the fruitful discussions on the beam dump design constraints and the magic of cyclotrons.Thisresearch wassupported bythetensiletests atReferences[1][2][3][4][5][6][7][8][9][10]T. T. Wohlers, I. Campbell, O. Diegel, R. Huff, J. Kowen, Wohlers Report 2019, Wohlers Associates Inc., March 2019. C. Y. Yap et al., ‘Review of selective laser melting: Materials and applications’, Applied Physics Reviews, vol. 2, no.

34, 80335 Munich, Germany *Corresponding author: maximilian.binder@igcv.fraunhofer.deAbstractThe production of complex lattice structures made of pure tungsten can be of great interest for potential applications in various industrial sectors such as energy technology or medical devices. One example is the plasma-facing armour of so-called limiter components in nuclear fusion power reactors, where the tungsten lattice armour is supposed to withstand extreme heat flux loads up-on transient plasma events.. The reliability of the tungsten armour is hence an important requirement for the sustainable operation of fusion power reactors [1,2]. Tungsten is difficult to process to a satisfactory degree due to its high melting point, its hardness as well as its susceptibility to cracking. Therefore, this paper presents the manner in which tungsten can be processed into fine lattice structures by means of high- temperature laser-based powder bed fusion.

Therefore, this paper presents the manner in which tungsten can be processed into fine lattice structures by means of high- temperature laser-based powder bed fusion. It also explains to what extent the used metal powder and the laser-exposure strategy have an influence on pores and component defects. It is shown how particle size distribution and sphericity of the powders have a major impact on the basic processability of the material. Furthermore, it presents to what extent the laser exposure parameters, such as the laser hatch distance, can have an influence on the resulting density of the material and which methods are used to determine the actual material density of lattice cubes in the first place.

Finally, measurements of the electrical conductivity of the fabricated AM structures are presented, as this is of interest with respect to many other areas of application.Keywords: Additive Manufacturing, High-Temperature Laser-based Powder Bed Fusion, Tungsten, Refractory MetalIntroductionDue to its high melting point, high thermal conductivity, high threshold energy for sputtering and low tritium solubility, tungsten has been the most favored candidate for the plasma-facing material (PFM) of future magnetic confinement fusion reactors [3]. The EU demonstration fusion reactor (DEMO), which is currently in a conceptual design phase, will be equipped with a tungsten wall [1,2,4]. Some of the tungsten-based plasma-facing661components require partly complex geometries to maximize their functionality.

Some of the tungsten-based plasma-facing661components require partly complex geometries to maximize their functionality. For instance, the tungsten armour of the plasma limiter components shall be designed in form of a fine lattice structure, which can hardly be achieved through conventional machining. This is one example why additive manufacturing of such parts is in high demand in this field of technology [3,5]. Additive manufacturing (AM) processes, such as Laser-based Powder Bed Fusion (PBF-LB), provide a unique freedom of design for developers. The AM process is an automated and tool-free process for producing three-dimensional physical components directly from a 3D CAD dataset [6,7]. However, AM also reaches its limits when it comes to processing special materials such as tungsten. For example, the brittleness and high ductile-to-brittle transition temperature (DBTT) of tungsten have a major impact on pores and cracks in the final parts [8].

For example, the brittleness and high ductile-to-brittle transition temperature (DBTT) of tungsten have a major impact on pores and cracks in the final parts [8]. Nevertheless, the formation of such defects can be reduced by using a high-temperature process. Still, a satisfactory part quality is difficult to achieve today, which points out the necessity of additional research [9]. Also, great effort has been put into optimizing numerous other aspects that can significantly influence the part quality, for example properties of the raw material or machine parameters used.Materials and MethodsThe following section presents the various methods and devices used for powder and part quality analysis, including the used PBF-LB/M facility and the design of experiments.Tungsten powder analysisFlowability The quality of a powder can be characterized by its flowability, which is influenced by many variables such as the particle morphology or particle size distribution.

Although it is not possible to describe the flowability as a function of these variables in general, there are special measuring instruments to determine values which can describe the flowability of powders [10]. To characterize the flowability, the following quantities were determined by using the FT4 powder rheometer by Freeman Technologies: Basic Flow Energy, Stability Index, Flow Rate Index, Specific Energy and Conditioned Bulk Density.The Basic Flow Energy (BFE) can be determined by measuring the work when moving a helix shaped blade downwards through a known volume of powder contained in a vessel. Depending on the type and condition of the powder, the blade experiences a certain resistance during the consolidation process. A low BFE, which results from low resistance, represents good flow properties of a powder. At the same time, a poor flowability is expected for high BFEs. The measurements can be made with non-variable and variable blade speeds [11].

At the same time, a poor flowability is expected for high BFEs. The measurements can be made with non-variable and variable blade speeds [11]. For the following measurements, a non-variable blade speed was chosen. The Stability Index (SI) is used to test whether a powder is affected by measuring. To determine the SI, a number of identical measurements using a constant blade speed are carried out in succession. In this study, seven measurements with a constant blade speed were conducted and the SI calculated as follows [12]:662𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸 𝑇𝑇𝐸𝐸𝑇𝑇𝑇𝑇 7 𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸 𝑇𝑇𝐸𝐸𝑇𝑇𝑇𝑇 1𝑆𝑆𝑆𝑆 =Energy Test 7 refers to the work done by the blade during the seventh downward traverse in mJ and Energy Test 1 during the first traverse, both with the same blade speed. An SI ≈ 1 indicates that the powder is not affected by the measurements. SI > 1 or SI < 1 indicates that a powder was affected during the measuring process.

An SI ≈ 1 indicates that the powder is not affected by the measurements. SI > 1 or SI < 1 indicates that a powder was affected during the measuring process. Possible reasons could be agglomeration or moisture uptake for SI > 1 and de-agglomeration or attrition for SI < 1 [12]. A powder that is moved around by a rake, for example, during an AM process will not behave the same way at all times. This is manifested by reduced flow rates in some areas and higher rates in other areas. The Flow Rate Index (FRI) provides information about how sensitive a powder is to flow rate. To determine the FRI, four additional measurements (measurements 8-11) were conducted, each by using a decreasing blade speed. In a second step, the results were compared as shown by the following equation [13]:𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸 𝑇𝑇𝐸𝐸𝑇𝑇𝑇𝑇 11 𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸 𝑇𝑇𝐸𝐸𝑇𝑇𝑇𝑇 8𝐹𝐹𝐹𝐹𝑆𝑆 =Energy Test 11 refers to the work done by the blade during the eleventh downward traverse in mJ and Energy Test 8 during the eighth.

Different blade speeds are used for test eleven and eight. Table 1 shows typical values for the FRI and their corresponding meanings. 1.5 < FRI < 3 is the average Flow Rate Index for most powders, whereas FRI > 3.0 indicates a high flow rate sensitivity and FRI < 1 a low sensitivity. FRI ≈ 1 signifies that the powder is flow rate insensitive.Table 1: FRI rangeFRI > 3 1.5 < FRI < 3 FRI ≈ 1FRI < 1Flow Rate Sensitivity high average insensitivepseudoplastic or NewtonianPowder very cohesive powders most powders large particle size or surface treatments powders containing flow enhancersThe Specific Energy (SE) is described as the energy per gram that is needed for the displacement when lifting a powder using a helix shaped blade. In contrast to the BFE, the powder is lifted upwards to produce shear and is not consolidated.

In contrast to the BFE, the powder is lifted upwards to produce shear and is not consolidated. This method is carried out to determine the flow resistance of powders without putting them under compression stress and to analyze the shear forces between particles. Thus, the SE can be an indicator for powder cohesion [14]. The Specific Energy is defined as:663𝑈𝑈𝑈𝑈 𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸 𝐶𝐶𝐸𝐸𝐶𝐶𝐶𝐶𝐸𝐸 6 + 𝑈𝑈𝑈𝑈 𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸𝐸 𝐶𝐶𝐸𝐸𝐶𝐶𝐶𝐶𝐸𝐸 7 2 ∗ 𝑆𝑆𝑈𝑈𝐶𝐶𝑆𝑆𝑇𝑇 𝑀𝑀𝑀𝑀𝑇𝑇𝑇𝑇𝑆𝑆𝐸𝐸 =Up Energy Cycle 6 refers to the energy measured during the upward movement of the sixth cycle and Up Energy Cycle 7 for the seventh cycle of a seven-cycle run, both in mJ. The split mass of the powder is given in g. Hence, the dimension of the SE is mJ/g. SE > 10 indicates high cohesion, values below 5 low cohesion, while 5 < SE < 10 indicates moderate cohesion [12].

The split mass of the powder is given in g. Hence, the dimension of the SE is mJ/g. SE > 10 indicates high cohesion, values below 5 low cohesion, while 5 < SE < 10 indicates moderate cohesion [12]. The Conditioned Bulk Density (CBD) is defined as the ratio between the split mass and the split volume of a considered powder, which is shown by the following equation:𝐶𝐶𝐶𝐶𝐶𝐶 =The dimension of the CBD is g/ml [12].𝑆𝑆𝑈𝑈𝐶𝐶𝑆𝑆𝑇𝑇 𝑀𝑀𝑀𝑀𝑇𝑇𝑇𝑇 𝑆𝑆𝑈𝑈𝐶𝐶𝑆𝑆𝑇𝑇 𝑉𝑉𝑉𝑉𝐶𝐶𝑉𝑉𝑉𝑉𝐸𝐸Particle Size Distribution (PSD) To determine the PSD, the Mastersizer 3000 by Malvern Panalytical and its corresponding software Mastersizer 3000 were used. A dispersion or suspension consisting of a powder particle collective and a liquid, e.g., demineralized water, is irradiated with laser light in a measurement cell. Different scattered light patterns, caused by diffraction, are detected depending on the size of the particles.

Different scattered light patterns, caused by diffraction, are detected depending on the size of the particles. Smaller particles diffract the laser light with low intensity and larger particles with high intensity. By using mathematical algorithms, the recorded diffraction and scattered light patterns are converted into particle size distributions. In this study, a suspension of five grams of each powder and 600 ml of demineralized water were used for the PSD measurements. Each powder suspension was measured three times for a period of ten seconds and the device was cleaned three times with demineralized water after each measurement.Microscopy A TM3030 Plus scanning electron microscope (SEM) by Hitachi and the TM3030Plus software were used to visualize the powder particle shapes of the relevant powders. For the EDX analysis, Quantax 70 software was chosen.PBF-LB/M using an AconityOneThe PBF-LB machine used for this study was the research unit AconityOne by Aconity GmbH.

For the EDX analysis, Quantax 70 software was chosen.PBF-LB/M using an AconityOneThe PBF-LB machine used for this study was the research unit AconityOne by Aconity GmbH. AconityOne is equipped with a 400 W laser (maximum power) with a beam diameter ranging between 80 - 500 µm. The PBF-LB unit reaches scan speeds up to 4 m/s and a minimal layer thickness of 10 µm. Furthermore, this machine provides a substrate heater creating substrate temperatures up to 800 °C. A residual oxygen content below 100 ppm can be reached by using argon as a protective gas, which floods the build space during processing [15].Part quality analysis664The fabricated grid cubes were separated from the building platform by wire cutting. The Archimedean method, according to DIN 66137-3 (2019), was applied to determine the part density.

The Archimedean method, according to DIN 66137-3 (2019), was applied to determine the part density. By comparing the weight of a specimen in air (dry weight) with the weight measured in a liquid (wet weight), this method allows the determination of the absolute density of a given part. Isopropyl alcohol was used to determine the wet weight in order to guarantee a complete coating of the specimen and a deep penetration of the liquid in the complex and finely shaped cubes. Each specimen was weighed three times in air and three times in isopropyl alcohol. Since the density of isopropyl alcohol at the given temperature is known, the part density was then calculated using the following equation:𝑉𝑉𝑎𝑎 ∗ 𝜌𝜌𝑙𝑙 (𝑇𝑇) 𝑉𝑉𝑎𝑎 − 𝑉𝑉𝑙𝑙𝜌𝜌 =where air, at a given temperature in g/cm³. The temperature of the liquid was measured before weighing. The density of the specimens was additionally determined by using the gas pycnometer Ultrapyc 1200e from Quantachrome.

The temperature of the liquid was measured before weighing. The density of the specimens was additionally determined by using the gas pycnometer Ultrapyc 1200e from Quantachrome. For this purpose, the density of at least 3 samples was simultaneously measured. Therefore, the specimens were weighed beforehand and then placed in the sample cell, which was flooded with helium. The volume of the gas inside the cell was measured and then compared to the volume of the gas in an empty sample cell, which allows the true density of the specimens to be determined. For the metallographic cross section analysis, some specimens were cut in half in the building direction using wire cutting. Subsequently, all samples were cleaned with ethanol by using an ultrasonic cleaning machine, which allowed surface impurities and excess powder to be removed from the parts. The specimen that were cut in half were cold embedded using KEM 60 embedding material from ATM GmbH.

The specimen that were cut in half were cold embedded using KEM 60 embedding material from ATM GmbH. Afterwards, the specimens were grinded and polished on a Saphir 530 machine from ATM GmbH. After two grinding processes, using water and diamond grinding plates called Galaxy Red and Galaxy Blue, the specimens were polished in three steps. Every specimen was cleaned using an ultrasonic cleaning machine between each of the steps to remove particles from the specimen surface. For polishing, increasingly finer plates and fabrics were used (Gamma, Zeta, Omega), together with Polycomplete Diamond suspensions for the steps of pre-polishing and polishing. A mixture of water and eposal was used for final polishing. Afterwards, the ground and polished specimens were carefully cleaned before the microscopic analysis. For the qualitative cross section analysis, the BX53M light microscope by Olympus and Stream Essentials software were used.

For the qualitative cross section analysis, the BX53M light microscope by Olympus and Stream Essentials software were used. The magnification was chosen so that relevant sections were clearly visible. While the light microscope is sufficient to make pores visible, the scanning electron microscope was used for more detailed investigations.is the part density in g/cm³,the mass in grams of the specimen measured inthe mass of the specimen measured in liquid and 𝜌𝜌the density of the liquid𝑉𝑉𝑎𝑎𝑉𝑉𝑙𝑙𝜌𝜌𝑙𝑙 (𝑇𝑇)MethodsIn a first step, the morphology of different tungsten powders is captured by using the scanning electron microscope. Then, the particle size distribution of the powders is determined, followed by the measurement of flow properties. After the selection of665specific powders, based on the powder analysis results, lattice structures are built by using the initial parameters from [16] shown in Table 2.

After the selection of665specific powders, based on the powder analysis results, lattice structures are built by using the initial parameters from [16] shown in Table 2. This includes a variation of the Laser Power and Scan Speed and the fabrication of 27 specimen. Afterwards, these lattice cubes are analyzed and an adapted parameter set is selected according to the cubes with the highest quality regarding part density. In the second step an additional parameter optimization, which includes the variation of the hatch distance and outer distance, is carried out to investigate the influence on the part density. For each build job, a tungsten substrate and a substrate temperature of 600 °C is chosen to produce the specimen.

In a final step, the electrical conductivity of the fabricated AM structures is measured by using the Sigmascope SMP350 eddy current meter from Fischer and the conductivity measuring device Loresta GX from NH Instruments.Table 2: Parameter optimization variations.Parameter Base from [16]: Laser Power in W: 400 Scan Speed in mm/s: 510 * Hatch Distance in µm: 80 Outer Distance in µm: 60 Layer Thickness in µm: 40 Energy Density in J/mm³: 245Parameter optimization 1: Variation of Laser Power and Scan SpeedParameter optimization 2: Variation of Hatch Distance and Outer DistanceLaser Power in W340 – 400400Scan Speed in mm/s500 – 700570 **Hatch Distance in µm8060 – 100Outer Distance in µm6060 – 100Layer Thickness in µm4040152 – 250Energy Density in J/mm³ * scan speed optimized for bulk material ** value adapted due to analysis results of filigree lattice structures175– 292Powder QualityEight powders from various powder manufacturers were examined as part of this study.

The images of the analyzed tungsten powders show predominantly round and even particles as seen in Fig. 1. However, irregularities such as dents are partially visible, which stem from incomplete spheroidization of the particles. Also, single particles of granular shape appear in the images of the powders. These are either also the result of incomplete spheroidization or they are foreign particles, which already existed or stem from contamination during the powder handling process. The EOS, H.C. Starck, HC4000 and TPT tungsten powders consist of non-spheroidized particles of irregular shapes and different sizes. The determined PSD curves are widely distributed (Fig. 2). The powder distributions are all monomodal and display their peaks in a range from 10 µm to 32 µm. Furthermore,666narrow and wide PSDs can be distinguished. Whereas the Tekna and EOS powder have the narrowest PSD, powders from LPW (UK82672 and UK81678) show a very broad PSD.Fig. 1.

Furthermore,666narrow and wide PSDs can be distinguished. Whereas the Tekna and EOS powder have the narrowest PSD, powders from LPW (UK82672 and UK81678) show a very broad PSD.Fig. 1. SEM images of the analyzed tungsten powders.35% 3025y t i s n e D e m u l o V2015105001020304050607080µm 90SizeclassUK82672UK81678UK80580TeknaHC4000TPTEOSFig. 2. PSD curves of the analyzed tungsten powders.667100GTPTable 3 summarizes the results of the flowability measurements. It shows the tendency of the Basic Flow Energy to increase for granular powders. More than two times less energy is required for the blade of the rheometer to move through the spherical Tekna powder compared to the granular EOS powder, which reaches values over 3000 mJ. The force between smaller particles is greater, which leads to higher BFEs. The BFE of the granular powders UK82672 from LPW and HC4000 from H.C.

The force between smaller particles is greater, which leads to higher BFEs. The BFE of the granular powders UK82672 from LPW and HC4000 from H.C. Starck show almost the same values as the EOS powder, reaching a maximum of 3477 mJ. The spherical powders range between 1250 mJ and nearly 1500 mJ. Partially spheroidized powders range around 2000 mJ. Regarding the Stability Index, most powders are settled at around SI = 1, indicating that the powders are not prone to measurement. However, the EOS powder seems to be affected by the measurements, with an SI = 1.2, assuming that possible reasons could be de-aeration, which is typical for powders with low bulk density. As air is squeezed out from the loosely packed powder, the density increases, causing more shearing between the particles, which could be the case for the HC4000 powder as well. Also, an agglomeration of particles during the measuring process can lead to an SI > 1 [12].

Also, an agglomeration of particles during the measuring process can lead to an SI > 1 [12]. The Flow Rate Index for each of the measured powders reaches values between 1.05 and 1.15. According to Table 1, these powders are unresponsive to variable external forces. Continuing with the Specific Energy, the results show a tendency for higher SEs for granular powders. Spherical and partially spherical powders clearly lie under 5 mJ/g, which is an indicator for low cohesion between the powder particles. The granular powders reach values between 4.5 and 5.3 mJ/g, which is in the spectrum for moderate cohesion. This is due to higher shear forces between particles of irregular shape [14]. The results for the Conditioned Bulk Density show that the spherical Tekna powder reaches the highest density of 11.85 g/ml. The UK82672 powder has the lowest bulk density of 8.44 g/ml, which is due to the granular particle shape and the consequently loosely packed powder bed.

The UK82672 powder has the lowest bulk density of 8.44 g/ml, which is due to the granular particle shape and the consequently loosely packed powder bed. The measurements for spherical powders result in about 25 % higher densities compared to granular powders. However, this trend should be critically viewed, as the CBD is also influenced by the PSD.Table 3: Results of the flowability measurements of the tungsten powders.Particle ShapeSphericalPartially sphericalGranularTungsten PowderUK80580 UK81678 Tekna TPT GTP HC4000 UK82672 EOSØ BFE in mJ 1282.67 1349.33 1460.33 1740.00 1922.00 3199.67 3477.67 3204.01Ø SI1.16 1.00 1.05 0.99 0.98 1.45 1.07 1.20Ø FRI1.17 1.06 1.18 1.07 1.07 1.14 1.11 1.15Ø SE1.72 2.02 1.99 2.08 2.90 4.92 5.29 4.55Ø CBD in g/ml 11.40 11.33 11.85 11.80 11.20 9.54 8.44 9.17Due to the high CBD, the Tekna powder is chosen to produce the following samples. Also, the rest of the measured values show convincing results for the Tekna powder.

Also, the rest of the measured values show convincing results for the Tekna powder. Furthermore, the EOS powder is chosen for comparison reasons.Resulting Build Quality for Lattice Structures668In the first optimization step, the laser power and the scanning speed were varied using the Tekna powder. Fig. 3 shows a CAD view and built version of the investigated lattice structure.A)B)z9 . 8 3 m myzxxC)zD)yyxFig. 3. Left: 3D view of lattice test specimen: A) isometric view including dimensions of the specimen, B) view of xz-plane, C) view of yz-plane, D) view of xy-plane. Right: Printed lattice-cubes on build plate.Parameter Optimization 1 Fig. 4 shows the relative part densities of the test specimens for Parameter Optimization 1. During the investigations, it was found that density measurement of the lattice structures via Archimedes’ principle becomes difficult due to the formation of air bubbles that are located in the numerous small cavities.

The EOS cubes have an average density of 86.38 %, with Cube 6 reaching the highest density of 86.99 % and Cube 7 the lowest value of 85.86 %.669Fig. 4. Density results of parameter variation Step 1 (laser power and scan speed using Tekna-powder). Combinations of laser power and scan speed that lead to the highest densities are highlighted.Fig. 5. Density results and comparison of lattice structures produced from Tekna and EOS powder (parameter: 340 W and 600 mm/s).In Fig. 6 the comparison of the results of the Pycnometer and Archimedean measurements is shown. For the Pycnometer, significantly higher values for the density of the Tekna and EOS lattice-cubes with a maximum difference of almost 6 % for the Tekna Cubes 1 – 4 and almost 9 % regarding the EOS Cubes 1 – 4 are measured. The highest average density is reached by Tekna Cubes 1 – 4 with a value of 97.14 %, whereas the lowest density is measured for the EOS Cubes 5 – 7 (94.39 %).

The highest average density is reached by Tekna Cubes 1 – 4 with a value of 97.14 %, whereas the lowest density is measured for the EOS Cubes 5 – 7 (94.39 %). The true density of the lattice structures is expected to be between the Archimedean and Pycnometer measurement results. This is because the measurements via Pycnometer670may not be able to detect existing defects, such as cracks or open pores, as defects, but rather assume them to be the outer surface of the component and thus a higher density is measured than is actually present.Fig. 6. Density results of lattice structures produced from Tekna and EOS powder. Comparison between density measurements using Archimedean method and Pycnometer (parameter: 340 W and 600 mm/s).An overview of cross section images of the tungsten cubes is shown in Fig. 7.

Comparison between density measurements using Archimedean method and Pycnometer (parameter: 340 W and 600 mm/s).An overview of cross section images of the tungsten cubes is shown in Fig. 7. The fact that the density of grid structures is lower than that of solid components manufactured with the same parameters may be due to the fact that grid structures conduct heat less effectively via the thin structures because they have less volume. This can cause a build- up of heat, resulting in pores and cracks due to excessive temperature gradients. Here (Fig. 7, F), it can be seen that a lack of fusion occurs between two melt tracks because the hatch distance is too large, resulting in a gap between two melt tracks. Therefore, the hatch distances have to be re-selected.671Fig. 7. Cross section images of Tekna- and EOS-powder lattice-cubes.

Therefore, the hatch distances have to be re-selected.671Fig. 7. Cross section images of Tekna- and EOS-powder lattice-cubes. A-C: SEM images of Tekna lattice-cubes cross sections with increasing magnification, D-F: xy-plane upskin SEM images of Tekna lattice-cubes with increasing magnification from left to right (no cross section). Nota bene: all specimen were fabricated on an Aconity One machineParameter optimization 2 In Parameter Variation 2, the hatch distance was investigated. The hatch distance is classified into the infill hatch distance and the outer distance as illustrated in Fig. 8 on the righthand side. The optimization was applied to minimize the porosity during the later limiter armor manufacturing. The following sample (Fig. 8 left) was built for the limiter manufacturing.

The optimization was applied to minimize the porosity during the later limiter armor manufacturing. The following sample (Fig. 8 left) was built for the limiter manufacturing. Since analysis and comparability of complex lattice structures was found to be difficult during Parameter Variation 1, a representative lattice plane (xy-plane) was selected in an isolated observation.672Fig. 8. View of the small-scale limiter mock-up: A) isometric view including dimensions of the specimen B) view of xy-plane C) view of xy-plane D) (Extraction of representative plane and illustration of outer distance and hatch distance with toolpath visualization).The results of the part density (measured via Archimedes) of the test series are shown in Fig. 9. The highest part density is produced at a combination of a hatch distance at 0.10 mm and outer distance at 0.10 mm.

9. Density results of Parameter Variation 2 (3 samples for each parameter variation).The results of the electrical conductivity measurements are shown in Fig. 10. For this purpose, the electrical conductivity of conventionally manufactured solid tungsten material was compared with additively manufactured tungsten structures (Tekna powder) by using the Lorasta GX and Sigmascope SMP350 devices. The results for the conventionally manufactured tungsten samples range between 16.56 and 18.81 MS/m, which is consistent with literature values (18.2 MS/m, [17]). In comparison, the measured values673for the additively manufactured samples are significantly lower. However, this can usually be compensated for in PBF-LB/M-materials by common post-treatments for the material. Sigmascope SMP350Loresta GX20MS/m 18161412Bulk material (conventionally manufactured)10Bulk material (additively manufactured)86420Fig. 10. Conductivity measurement results.

10. Conductivity measurement results. Left: Loresta GX measurement results. Right: Sigmascope SMP350 measurement results.DiscussionAs a conclusion of this investigation, it can be noted that the optimum process parameters strongly depend on the geometry of the manufactured components (solid vs. thin walls vs. grids). Accordingly, the parameters have to be adjusted when the geometry is changed. Parameter optimization on grids exceeds standard analysis methods. They are no longer applicable to the same extent. For an improvement of the part density, it would make sense to define the laser vectors manually for a unit cell. As the investigations during the parameter optimization in the second variation showed, the arrangement of the hatches plays a significant role in the complex lattice structures. The automatic arrangement of the laser vectors by the slicing software does not lead to the desired high densities.

The automatic arrangement of the laser vectors by the slicing software does not lead to the desired high densities. Therefore, a further investigation would require individually fitting laser vectors (and hatch distances) for the second tested geometry. These would then have to be patterned so that whole limiter mock-ups could be produced with high densities.Furthermore, an adaptation for the determination of the density via Archimedes was tested for the analysis of tungsten grid structures: a procedure is developed which reduces the number of gas bubbles when measuring the density of the tungsten grids. Thus, the samples are to be placed, as usual, in isopropyl alcohol for the determination of the density. Before the sample is weighed, ultrasound is applied to the container carrying the liquid and the sample. This induces the creation of air bubbles, which adhere in the grid.

Before the sample is weighed, ultrasound is applied to the container carrying the liquid and the sample. This induces the creation of air bubbles, which adhere in the grid. After an appropriate time, when no more formation of air bubbles from the sample can be noted, the measurement is carried out using Archimedes’ principle (without taking the sample out of the liquid). The sample is dried for several days before it is subsequently674weighed dry. For this purpose, ten additional lattice-cubes were produced by using the Tekna powder and the optimized parameters from Variation 2. In a first step, the cubes’ densities were measured as usual using the Archimedes’ principle (without ultrasound treatment). Afterwards, the above mentioned adaption of the procedure was applied to determine the density. The measurement results in Fig. 11 show noticeably higher values compared to the prior results. In case of Cube No.

11. Comparison of the density measurement results via Archimedean method with and without ultrasound treatment of the lattice-cubesSummary and OutlookIn this study it was shown that lattice-cubes fabricated by using an Aconity One machine and granular powder clearly result in lower part densities compared to the cubes produced on the same machine but from spherical powder. By using the same parameters the fabrication of the granular powder resulted in more porous parts compared to spherical powders, which corresponds with the part density values. Thus, the selection of the raw material plays a crucial role regarding the part quality. However, in the additive manufacturing of complex, thin lattice structures, the main challenge is the large number of different layers.

However, in the additive manufacturing of complex, thin lattice structures, the main challenge is the large number of different layers. In these layers the wall thicknesses of the struts, the degree of overhanging areas or the proportion of solid material varies considerably, which is why lattice-structures show significantly lower densities compared to bulk material. Thus, if a parameter set is to be determined that provides the best possible result for all these areas, it is often feasible for materials that can be welded well, but very challenging for tungsten structures or other refractory metals.675In the study conducted by Müller et al. [16] a parameter set was developed that is especially well suited for the fabrication of honeycomb structures and solid material areas. In this framework, the laser hatches of the honeycomb structures were designed and generated manually, which represents a significant human effort.

In this framework, the laser hatches of the honeycomb structures were designed and generated manually, which represents a significant human effort. This ensured that the laser exposure vectors met the requirements better and that a higher material density could be generated with a lower number of cracks and pores. For the fabrication of lattices, which are much more complex than the geometry of honeycomb structures, the laser hatches were automatically generated using software, as is common in additive manufacturing. This creates areas in the lattices that are exposed with less or more than the desired hatch distance. Through this study, it now stands to reason that this very factor is responsible for a higher number of pores and cracks in the structures than usual. Thus, with a unit cell height of 1.52 mm and a layer thickness of 0.04 mm, 38 layers (number of layers = height / layer thickness) would have to be adjusted in their laser-hatch arrangement.

Thus, with a unit cell height of 1.52 mm and a layer thickness of 0.04 mm, 38 layers (number of layers = height / layer thickness) would have to be adjusted in their laser-hatch arrangement. In addition, it is a major finding that the currently available software programs are not sufficient for optimal processing of tungsten. In the further course of research, a three-stage approach would be useful in order to guarantee the highest possible quality of future manufactured samples and products:1) Manual adjustment of laser hatches for the lattice cube geometry to verify the established thesis.

The cube can then be manufactured with satisfactory quality (e.g., density > 98.5 %).2) Derivation of hard, geometry-dependent criteria, for reliable manufacturing of for (e.g.,3) Development and implementation of a suitable software solution to transfer thefindings to any other type of lattice structures.Finally, it was shown that an adaptation for the determination of the density for lattice structures via Archimedes by removing air bubbles and access powder through ultrasound is more accurate compared to the standard procedure.

lid Freeform Fabrication 2017: Proceedings of the 28th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference Reviewed PaperInfluence of process time and geometry on part quality of low temperature laser sinteringYuki Yamauchi*, Toshiki Niino†, and Takashi Kigure*Tokyo Metropolitan Industrial Technology Research Institute †Institute of Industrial Science, the University of TokyoAbstractThe authors are developing a novel laser sintering process that prevents parts from warping by anchoring them to a rigid base plate. Since the powder bed temperature of the process is normally lower than in the standard process, the laser is required to supply more energy in the novel process, namely low temperature process. Accordingly, the part quality is more sensitive to laser parameters. Additionally, accumulation and dispersion of energy which is supplied by the laser through layers plays an important role in the consolidation of the powder.

Additionally, accumulation and dispersion of energy which is supplied by the laser through layers plays an important role in the consolidation of the powder. Thus, in low temperature process, parameter relating part geometry and time affects the part quality more than in standard high temperature process. In this research, the influence of part size and process time per layer on the density of parts as a primary index of part quality is investigated. Density decreases as the process time per layer increases. With respect to part size, density increases as parts become larger.IntroductionAmong various AM technologies, plastic laser sintering (LS) is promising for end parts production. Although LS technology has been utilized in aerospace and medical industries for more than ten years, expansion into other industries is still limited. One of the greatest drags on wide spread of the technology is material cost. There are two causes that elevate the cost.

One of the greatest drags on wide spread of the technology is material cost. There are two causes that elevate the cost. One is the high price for which we purchase powder. Normally, the price of LS powder is much higher than pellets for injection molding due to the cost of producing powder from raw material and initial development. The other cause is the large amount of powder we use in the LS process. When an LS process is finished, parts are buried in powder that remains un-consolidated. The powder is not fused intentionally but slightly fused to make a very porous cake, known as part-cake. Usually, the part-cake is separated from the parts, sieved to obtain powder form again, and reused in the next batch. Before it is used again, a certain amount of brand-new powder is added to reduce the effect of the previous process. The effect is supposed to be deterioration of polymer that is caused by preheating [1].

The effect is supposed to be deterioration of polymer that is caused by preheating [1]. The maximum value in rate of used powder in feed stock that enables a successful process is defined as “recycle rate” or “recyclability.” Low recyclability directly impacts material cost of part production using laser sintering. In practical LS operation, the ratio of parts to whole bed is 5 to 10 %. On the other hand, recyclability of commercially available PA12 powder is 50 to 70 %, and the amount of the additional fresh powder is 30 to 50%. Thus, the amount of powder that we have to purchase can be ten times as much as the powder that had been consolidated to parts. It is supposed that the reduction of the deterioration improves the recyclability and reduces the cost, consequently.

It is supposed that the reduction of the deterioration improves the recyclability and reduces the cost, consequently. Additionally, it is reported that the preheat temperature is one of the dominant factors on powder deterioration [2, 3, 4].The authors proposed a novel LS process in which part warpage is suppressed by tying parts to a rigid base plate during the process instead of high temperature preheating [5, 6]. This process allows the powder bed temperature to be lower than recrystallization temperature.

This process allows the powder bed temperature to be lower than recrystallization temperature. In the following description, we denote the novel process as “low temperature process.” Contrarily, the normal1495process in which powder bed temperature is maintained above recrystallization point is denoted as “high temperature process.” Although the low temperature process can provide better recyclability, as mentioned, part quality out of the process is not investigated in details.When the bed temperature is lower, the laser is required to supply more energy than in high temperature process [7]. Accordingly, part quality is more sensitive to laser scan parameters, part geometry, and parts arrangement in low temperature process than high temperature one. The aim of this research is to reveal dominant factors that affect part quality in low temperature process.

The aim of this research is to reveal dominant factors that affect part quality in low temperature process. Previously, it was reported that mechanical properties is very much depending on geometry and time to process one layer in high temperature process [8, 9, 10, 11]. In this paper, we focused on process time that is required to process one layer, hereinafter referred to as layer time, and part geometry as parameters of accumulation and dispersion of supplied heat energy.Material and MethodsLaser sintering apparatusA commercially available LS machine (RaFaEl 300F, Aspect Inc.) was used. This machine is installed with a fiber laser while the typical commercially available machine uses a CO2 laser. Since the fiber laser has a shorter wave length than the CO2 laser, its beam is focused into a small spot more easily.

Since the fiber laser has a shorter wave length than the CO2 laser, its beam is focused into a small spot more easily. Basic specifications of this machine are summarized in Tbl.1.Tbl.1 Specifications of LS machineSystemRaFaEl 300FType of laserFiber laser (λ=1064nm)Nominal laser spot diameter170 μmMaximum laser power20.0 WWork volume290 mm×290 mm×370 mmMaterialPA11 powder (Aspex-FPA, Aspect Inc.) was employed. Since light from the fiber laser is not absorbed by the base resin of this powder, black pigment dedicated to improve absorption is compounded in each powder grain (Fig.1). Properties of this powder are summarized in Tbl.2. This material recrystallizes in the rage between 160 °C and 175 °C.

Properties of this powder are summarized in Tbl.2. This material recrystallizes in the rage between 160 °C and 175 °C. In low temperature process, the powder bed is preheated at 140°C, which is lower than the recrystallizing range.1496200μmFig.1 Optical micrograph of PA11 grains Tbl.2 Material properties of Aspex-FPAType of materialPA11Melting point201 °CRecrystallization temperature167 °CTrue density1.01 g/cm3Average particle size50 μmBase Plate for low temperature processThe base plate to which parts were anchored to prevent warpage during process was prepared. It was a flat PA12 (Aspex-PA, Aspect Inc.) plate of 13mm×200mm×200mm as shown in Fig. 2. The plate was manufactured using high temperature process with a commercially available LS machine (RaFael 550, Aspect Inc.).

2. The plate was manufactured using high temperature process with a commercially available LS machine (RaFael 550, Aspect Inc.). The plate was fixed to a back-up plate with screws to ensure the flatness of the anchoring surface.Fig.2 Schematic view of base plate1497Density measurementDensity of sintered parts has a strong correlation with mechanical properties [12]. of aTherefore, relative density of specimen was evaluated as a part quality. Relative density specimen can be obtained by𝜌𝜌𝑟𝑟𝑚𝑚 are mass of the part, part volume and true density of the powder material. Here, 𝑣𝑣𝑏𝑏 × 𝜌𝜌𝑡𝑡 of the material employed in this research is 1.01 g/cm3. The mass of part was measured by an electronic balance (AUX220, Shimadzu Inc.). The part volume was calculated by measuring the 𝜌𝜌𝑡𝑡 dimensions of each axis direction with calipers and micrometers.where,and𝜌𝜌𝑟𝑟 =𝑣𝑣𝑏𝑏𝜌𝜌𝑡𝑡𝑚𝑚Build parameters of each processThe build parameters are summarized in Tbl.3.

For comparison, a specimen was also built with the high temperature process using the parameter set that is recommended by machine and material manufacturer. Parameters for low temperature process was decided so that obtained density becomes equivalent to that from high temperature process [13]. Laser was scanned in the direction of x-axis, and layering direction was z-axis.Tbl.3 Build parametersProcessLaser powerScan speedScan intervalEnergy per unit areaLayer thicknessPowder bed temperatureLow temperature High temperature15W10W2.0 m/s10.0 m/s90 μm90 μm83 kJ/m211 kJ/m2100 μm100 μm140 °C190 °CRelationship between layer time and density of partThe influence of layer time on the density of the part was tested. Here, we define the layer time ( = ) as the time between two layering operations. Specimens were built in each layer time ( 5s, 15s, 25s, 35s, and 45s) and the density of the obtained parts was measured.

Specimens were built in each layer time ( 5s, 15s, 25s, 35s, and 45s) and the density of the obtained parts was measured. The layer time consists of powder spreading, laser scanning, and platform descending. To set the layer time to specific value, a pause is inserted in the process and the pausing time was adjusted. Specimens that were designed in 20mm×10mm×5mm were built.𝑡𝑡𝑙𝑙The influence of the sequence of processing in a layer was investigated as well. Separated regions as shown in Fig. 3 were processed in the order indicated by the number of each region. Density of hatched parts was measured to evaluate the effect of process order.1498𝑡𝑡𝑙𝑙Fig.3 Arrangement of specimens and process sequenceRelationship between part geometry and densityThe effect of changing size and aspect ratio of parts were investigated. Specimens that are the same in aspect ratio of cross section and different in area as shown in Fig.

Specimens that are the same in aspect ratio of cross section and different in area as shown in Fig. 4 are fabricated, and their densities were measured to evaluate how the difference in cross section area effects the part property.𝐴𝐴𝑥𝑥𝑥𝑥Additionally, specimens that are the same in cross sectional area and different in aspectratio, in contrast, as shown in Fig. 5 were tested as well.Fig.4 Cross-section area of specimens with same aspect ratio (thickness: 5mm)1499Fig.5 Aspect ratio of specimens with same cross-sectional area (thickness: 5mm)Experimental resultRelationship between layer time and density of partA relationship between layer time and relative density of specimens is shown in Fig.6. In the high temperature process, the variation of relative density was smaller than 3%. As a general decreased as the layer time increased in the low temperature process.

In the high temperature process, the variation of relative density was smaller than 3%. As a general decreased as the layer time increased in the low temperature process. Part density tendency, decreased as layer time increased, and the decrease stopped when the layer time was more than 35s. Relationship between order of processing in layer and relative density of specimens was shown in Fig.7. The difference in each order of processing on relative density was smaller than 2%. When the part number was bigger (that is to say order of processing was the latter), the relative density slightly decreased. When specimens shown in Fig.3 are processed, the layer time was 30s.𝜌𝜌𝑟𝑟10095]%[ y t i s n e d e v i t a l e R90858075Powder bed temp. 140Powder bed temp.

When specimens shown in Fig.3 are processed, the layer time was 30s.𝜌𝜌𝑟𝑟10095]%[ y t i s n e d e v i t a l e R90858075Powder bed temp. 140Powder bed temp. 190℃7001020 30 Layer time [s]40℃50Fig.6 Relationship between layer time and relative density in each process150010095]%[ y t i s n e d e v i t a l e R90858075700 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 Part numberFig.7 Relationship between order of processing in layer and relative density in the low temperature processRelationship between part geometry and densityThe relationship between the cross-sectional area of the specimens and the relative density= 50 mm2, 200 mm2 and 450mm2,is shown in Fig.8. In the high temperature process, when = 800mm2, were 85%. When On the other hand, as a general tendency,was slightly decreased. 𝐴𝐴𝑥𝑥𝑥𝑥 = 800 mm2,𝜌𝜌𝑟𝑟 increased in low temperatureincreased as𝐴𝐴𝑥𝑥𝑥𝑥 = 50 mm2,𝜌𝜌𝑟𝑟 was 76%. And whenprocess.

When On the other hand, as a general tendency,was slightly decreased. 𝐴𝐴𝑥𝑥𝑥𝑥 = 800 mm2,𝜌𝜌𝑟𝑟 increased in low temperatureincreased as𝐴𝐴𝑥𝑥𝑥𝑥 = 50 mm2,𝜌𝜌𝑟𝑟 was 76%. And whenprocess. Whenwas 82%.𝜌𝜌𝑟𝑟𝐴𝐴𝑥𝑥𝑥𝑥The relationship between the aspect ratio of specimen and the relative density is shown in decreased as the aspect ratio increased. The relative density𝐴𝐴𝑥𝑥𝑥𝑥 Fig.9. In low temperature process, at a ratio of 50 is smaller than at a ratio of 1 by 5%.𝜌𝜌𝑟𝑟𝐴𝐴𝑥𝑥𝑥𝑥𝜌𝜌𝑟𝑟𝜌𝜌𝑟𝑟Fig.10 shows an X-ray CT image of horizontal cross-sections of specimens which were obtained by low temperature process. Brightness shows local density. The black (darker) region is air or a pore, and gray region is plastic. In the low temperature process, the density was uniform all over the section. Fig.11 shows a CT image of high temperature processed parts.

In the low temperature process, the density was uniform all over the section. Fig.11 shows a CT image of high temperature processed parts. Specimens out of high temperature process were relatively porous in the middle and dense at both ends in x direction or scanning direction. The width of the denser region was roughly 1 mm.Layer times for low and high temperature process in this experiment were 54s and 26s,respectively.15011009590]%[ y t i s n e d e v i t a l e R85807570Powder bed temp. 14065Powder bed temp. 190℃60℃02004006008001000Cross sectional area [mm2]Fig.8 Relationship between cross-sectional area of rectangular specimens with the same aspect ratio and their relative density10095]90%[ y t i s n e d e v i t a l e R85807570Powder bed temp. 14065Powder bed temp.

14065Powder bed temp. 190℃60110 Aspect ratio of specimen in horizontal section℃100Fig.9 Relationship between aspect ratio in horizontal cross-section of specimenswith= 200 mm2 and relative density𝐴𝐴𝑥𝑥𝑥𝑥1502Fig.10 X-ray CT image of horizontal cross-sections of specimens which were obtained by low temperature process (z = 2.5mm). Brightness shows local density.Fig.11 X-ray CT image of horizontal cross-sections of specimens which were obtained by high temperature process (z = 2.5mm). Brightness shows local density1503DiscussionInfluence of process time per layer on part qualityThe density of part became higher as the layer time became shorter. The effect of changing layer time is stronger when the time is relative low. This tendency can be explained as following. In low temperature process, the greater part of energy for melting is supplied by the laser.

This tendency can be explained as following. In low temperature process, the greater part of energy for melting is supplied by the laser. The supplied heat for each layer stays in the scanned region and contributes to the process of the next layer also. Since the heat disperses until heat for the next layer is supplied, a long layer time decreases the contribution of previously supplied heat. And the density of the parts decreases.Contrarily, in the high temperature process the powder bed is kept near melting point evenif the laser has not been irradiated, and, as a result, the contribution of process is relatively low.Influence of geometry on part qualityIt was found that the density of the part was very much dependent on its geometry. This can be explained as follows. When powder is melted, molten plastic flows and fills the gaps between powder grains.

This can be explained as follows. When powder is melted, molten plastic flows and fills the gaps between powder grains. The molten plastic needs a period of time to fill; temperature of plastic is required to be higher than the melting point for the period. In the low temperature process, the supplied energy and heat is dispersed to the surrounding powder. Since the dispersing rate is faster when parts are smaller, the density of low specimens are small. The dispersing speed also increases as the aspect ratio increases. Thus, the greater the aspect ratio is, the lower is the density.ConclusionThe influence of geometry and layer time on part quality in low temperature laser sintering was tested. As general tendency in this study, low temperature laser sintering is affected by layer time and geometry in the horizontal cross-section. The density of part is higher when the layer time is shorter in the low temperature process.

The density of part is higher when the layer time is shorter in the low temperature process. When the layer time was 5s, the density of the part was 95%, but when it was 45s, the density decreased until 82%. The influence of layer time is stronger when it is shorter, and becomes drastically small after 25s. On the other hand, the influence of the processing order in a layer is negligibly small. With respect to the part geometry, the density of part increases as the horizontal cross-section area of part increases, and the maximum difference was 6% in this paper.

With respect to the part geometry, the density of part increases as the horizontal cross-section area of part increases, and the maximum difference was 6% in this paper. The density of part decreases as the aspect ratio of the part in cross-section increases.On the other hand, in high temperature process, the density is affected by the part geometrywhen the scan length is longer.AcknowledgmentThis research is supported by “Strategic Innovation Promotion Program” Cabinet Office,Government of Japan and New Energy and Industrial Technology Development Organization.References[1] Beaman, et al, “Solid Freeform Fabrication: A New Direction in Manufacturing,” (1996) [2] Dotchev et al, “Recycling of polyamide 12 based powders in the laser sintering process,”Rapid Prototyping Journal, 15 (2009) 192-2031504[3] Kruth et al, “Consolidation of Polymer Powders by Selective Laser Sintering, ” Proc.

FLUENCE OF RHEOLOGY ON DEPOSITION BEHAVIOR OF CERAMIC PASTES IN DIRECT FABRICATION SYSTEMSBruceH. King, Sherry L. Morissette, Hugh Denham, Joseph Cesarano, III and Duane Dimos Sandia National Laboratories Albuquerque, NM 87185ABSTRACTRheology and deposition behavior of four commercially available thick-film inks and an aqueous alumina slurry were investigated using two different slurry-based deposition systems. The first of these deposition systems, a Micropen, is a commercially available system designed for the deposition of electronic thick film circuits. The second system, referred to as a Robocaster, is a developmental system designed to build thick or structural parts. Slurry rheology was seen to have a minor effect on deposition behavior and the bead shape when deposited using the Micropen. The deposition behavior was instead dominated by drying rate; too rapid of a drying rate led to excessive clogging of the tip.

The deposition behavior was instead dominated by drying rate; too rapid of a drying rate led to excessive clogging of the tip. Slurry rheology had a greater impact on the shape of beads deposited using the Robocaster. Highly viscous slurries yielded initially well-defined beads, whereas beads deposited using fluid slurries spread quickly. In both cases, significant spreading occurred with time. These observations only held for slurries with slow drying rates. It was observed that very fluid slurries produced well-defined beads when the drying rate was suitably high.INTRODUCTIONSlurry-based direct fabrication systems offer a number of advantages over conventional fabrication techniques and have capabilities not found on other direct fabrication systems. The material systems are relatively simple and a wide range of materials 'can be used, including pastes of metals, polymers or ceramics.

The material systems are relatively simple and a wide range of materials 'can be used, including pastes of metals, polymers or ceramics. Additionally, there are a number of commercially available pastes available from the screen-printing industry. Unique capabilities include multi materials/graded parts, incorporation of tailored porosity via fugitive materials and rapid turnaround of finished pieces. It is also possible to fabricate multilayer thick film circuits directly on a substrate, a difficult task using conventional screen printing techniques.Weare working with two slurry-based deposition systems at Sandia. The first is a commercially available system called Micropen.* The second is a system under development at Sandia called Robocasting. The Micropen is designed for use in the electronics industry for the deposition of thick film circuits and typically uses commercially available thick film inks.

The Micropen is designed for use in the electronics industry for the deposition of thick film circuits and typically uses commercially available thick film inks. l The design focus of the Robocaster is for thick or structural parts and typically uses aqueous slurries.2A schematic of the Micropen system is shown. in Figure 1, along with a circuit printed onto a curved surface. The system consists of a cantilevered tube to the end of which a pen tip· and permanent magnet are attached. An electromagnetic actuator positioned above the permanent magnet controls the height of the pen tip by varying the strength of the magnetic field. A height detector consisting of a flag attached to the tube, an LED and photodiode providesOhrrtcraft, Inc. Honeoye Falls, NY 14472391height feedback to the processor controlling the electromagnetic actuator. This system employs a stationary Z-axis; a pattern is written by pumping the slurry through the tube and moving an X Y table beneath the tip.

This system employs a stationary Z-axis; a pattern is written by pumping the slurry through the tube and moving an X Y table beneath the tip. A circuit is deposited as follows; the pen tip is brought into contact with the substrate and ink is pumped through the pen tube into the pen tip. As ink begins to build up under the pen tip, the pen tip rises until it reaches a predetermined trigger height. At this point, the .X-Y t~~le be?ins to move and the.pattern is deposited. A small downward force is applied durmg wrItmg VIa the electromagnetIc actuator. This helps to ensure that a uniform line of material is deposited. An advantage of this type of system is that in addition to the feedback provided by the height sensor, there is a force feedback provided by the electromagnetic actuator. Thus, the system can vary the pen height based on the force feedback to account for topography on the substrate.

Thus, the system can vary the pen height based on the force feedback to account for topography on the substrate. This enables a uniformly thick circuit to be deposited onto a curved surface or over existing features on a flat substrate. The Micropen typically uses pen tips in the size range of 25-250 /-lm. A typical slurry composition is 25-30vol% solids, 40-50vol% resin and 25 30vol% solvent.LED and photodiode height detectorElectromagnet (actuator)Permanent MagnetCantilever spring pentubePen TipForce fTriggerillJIlJ'JIlFigure 1: Schematic of the. Micropen system and a circuit printed on a curved surface.In contrast, the Robocaster;shown in.Figure 2, is a more straightforward design. The slurry is extruded .from a syringe mounted ona Z-axis above a moving X-Y table. Once a layer is completed, the Z-axis· is incremented. byapredetermined layer height and the next layer is deposited.

Once a layer is completed, the Z-axis· is incremented. byapredetermined layer height and the next layer is deposited. Currently, this system provides no fee<ibackabou.Uhe <actuaLdistance between the buildsurface.andthe pentip. However, a laser height systemis under devel9pment at Sandia that wiHprovi<iethis information. Thissystelll should allow for feedbacktothepumping system to account .forimpet"fections in the build ofprevious layers... The Robocaster typically uses. pen tips in the size range of 250-1500 /-lm. A typicaLslurry composition is very highly loaded with 58-61vol% solids, 40vol% water and a very small amount of processing aids such as Darvan and citric acid. Typical parts are also shown in Figure 2.Both systems rely on good bead definition to bu.ild accurate parts.

Typical parts are also shown in Figure 2.Both systems rely on good bead definition to bu.ild accurate parts. The electrical properties of electronic components deposited with the Micropen often depend on the accuracy of the line width which is in turn influenced by the degree of bead spreading after the line has been printed. Minimum feature sizes and the minimum spacing between· components are also392influenced by bead spreading. These aspects are important for both systems. Electronic components that are too close together may be subject·to ·shorting or electromigration. The dimensional accuracy of 3-D parts fabricated with the Robocaster is highly dependenton bead spreading. Additionally, significant bead spreading pfoneJayer leads to problems during the deposition of subsequent layers, resulting in gross defects within the finished piece. The rheology of theislurriesusedin these systems was investigated to gain an understanding of its influenc.e on bead shape.

The rheology of theislurriesusedin these systems was investigated to gain an understanding of its influenc.e on bead shape. This inf0I'Illation ",ill drive both system development and materials processing as new slurries are developed for each system.Figure 2: Schematic ofthe Robocasterand examples of complex shapes fabricated with it.EXPERIMENTALA number of slurries were investigated, four commercial thick filmpastesand.anaqueous alumina slurry prepared at Sandia. Two of the commercialinks were DuPont· posh~re inks; 5715, a gold conductor paste,and 1731, a ruthenium oxide-based resistor paste. /Bothofthese pastes are designed to be printed.·on dense alumina substrates and fired at high temp~r~tures.• (~ 1200°C). The othertwocommercial pastes werepolymerthic~ film (PTF>ipastes, Minicot M 4100.silver conductor paste and Asahi TU-1 k carbon resistor Paste.

The othertwocommercial pastes werepolymerthic~ film (PTF>ipastes, Minicot M 4100.silver conductor paste and Asahi TU-1 k carbon resistor Paste. Both ofthesepast~sconsist of solids (either silver or carbon) loade~ in~n epoxy carrier and are designed to be printed on FR4.epoxy board.• The epoxy carrier is cured at low temperatur~("'200°C).The(aqueous alumina slurry was prepared at Sandial)yball milling 60vol%AlcoaA-15aluminafor2 weeks in deionizedwat~r with Darvan 821A dispersantand citric acid.DuPont Electronic Materials, Research Triangle Park, NC 27709 tEmetson and Cuming, Lexingtotl,MA 02173 : Multicore Solders,Inc., Richardson, TX75081393The rheology of the five pastes was characterized using a controlled stress rheometer·. The commercial pastes were characterized with a 10 cone and plate geometry, while the aqueous alumina slurry was characterized using a cup and bob geometry. All rheology data was plotted as viscosity (cP) vs.

All rheology data was plotted as viscosity (cP) vs. shear rate (lis) on a semi-log plot.In addition to rheology, the slumping behavior of beads printed from each paste was investigated when deposited from each of the two systems. In all cases, a single bead of material was deposited onto the substrate and the shape of the bead was characterized via laser profilometry.t The post-fire inks were deposited onto alumina substrates, while the PTF inks were deposited onto FR-4 epoxy board. The aqueous alumina slurry was deposited onto dense alumina substrates. Pastes deposited with the Micropen were deposited using a 250 J.1m (10 mil) tip and bead profiles were obtained at times ranging from 1 minute to 1 hour after printing. Pastes deposited with the Robocaster were deposited with an 890 J.1m (35 mil) tip. In this case, n the equipment and profiles were obtained within"'" 10 seconds the profilometer was moun bead was performed for up to 30 minutes after printing. To of printing.

In this case, n the equipment and profiles were obtained within"'" 10 seconds the profilometer was moun bead was performed for up to 30 minutes after printing. To of printing. investigate the effect e substrate has on the profile, these experiments were repeated for the silver PTF paste deposite on a dense alumina substrate. From these profiles, the bead width and area as a function of time after printing were obtained.51054105 =- 3105 ~ ""= 2105 CJ;;CJ•D •o <) ... EElAu Conductor RuO Resistor Ag FrF Conductor Carbon PTF Resistor Aqueous AluminaPrint RangeFigure 3: Viscosity as a function of shear rate for commercial screen printing inks and aqueous alumina slurry.RESULTSViscosity as a function of shear rate for each of the pastes is shown in Figure 3. Included on the plot is the calculated range of shear rates experienced during printing? Of the slurries investigated, the post-firing inks were by far the most viscous.

Included on the plot is the calculated range of shear rates experienced during printing? Of the slurries investigated, the post-firing inks were by far the most viscous. At low shear rates, the PTF inks were significantly lower. The carbon resistor ink was the more viscous of the two, but they converged in the print range. At low shear rates, the aqueous alumina slurry exhibited a higherBohlin C8-10, Bohlin Instruments, Cranbury, NJ 08572 t Cyber8can Cobra, CyberOptics, Minneapolis, MN 55416394viscosity than the silver PTF conductor, but at high shear rates its viscosity was the lowest of the five pastes.

All pastes were observed to converge to the same viscosity at very high shear rates, however this is above the print range.15 ~-=-~-:----------,RuO2 Resistor1 Minute~ ~45MlB""25Figure 4: Profile of a resistor paste beadprinted with the Micropen at times of I minute and 45 minutes after printing.Bead SpreadingBead Sbrinkage20-~o oI10o ¢ 0 J;J !tl ~ 00 ~- -_ Targc:rr..Widt~Au Conductor 0 ¢ ...RuO2 Resistor Ag PTF Conductor Carbon PTF ResistorIIII30 Elapsed Time, min204050[6040 .r- 35 ==! 30 ~ ~ 25 'i= 20 ~PE.t:h-. v¢ ........ Ii II I - ... - ..I¢¢¢¢¢¢¢¢u 15110ttu8°ltdu~tor u . eSlstor~ ...

v¢ ........ Ii II I - ... - ..I¢¢¢¢¢¢¢¢u 15110ttu8°ltdu~tor u . eSlstor~ ... ¢Ag YrF Conductor Carbon PTF Resistor 1U 5 o oIIII2030 Elapsed Time, min104050Figure 5: Spreading and Shrinkage as a function of time after printing of thick film inks printed with the Micropen.Profiles similar to the profile shown for the ruthenium oxide resistor paste in Figure 4 were obtained for all of the pastes printed in the Micropen, with the exception of the alumina slurry. This slurry proved to be impossible to deposit with the Micropen due to excessive drying. Figure 5 plots width and cross-sectional area as a function of time after printing. Of the remaining four slurries, minimal spreading was observed after the first minute after printing. However, a significant change in the cross-sectional area of each bead was noted over the course of an hour.

However, a significant change in the cross-sectional area of each bead was noted over the course of an hour. This change in area may be attributed to drying shrinkage.In this case, only three pastes were printed for comparison; the highly viscous ruthenium oxide resistor paste, the relatively fluid silver PTF conductor paste and the aqueous alumina slurry. Initially, the ruthenium oxide resistor paste had a well-defined profile. However, the bead spread quiteFigure 6 compares the profiles of pastes printed with the Robocaster.395~~>60rapidly and continued to spread for up to 20 minutes. This is shown clearly in Figure 7, which plots bead width as a function of time after printing. The fluid silver PTF paste, on the other hand, spread significantly before the first profile was taken. After this, spreading continued but In contrast, the alumina slurry exhibited a well-defined bead initially. at a much slower rate.

After this, spreading continued but In contrast, the alumina slurry exhibited a well-defined bead initially. at a much slower rate. Unlike either of the thick film pastes, it exhibited no spreading with time. This is contrary to what would be expected based on the rheology measurements; the viscosity of the aqueous system indicated that it should have an initial profile similar to the silver PTF conductor paste and should probably have spread in a similar manner. However, as is shown in Figure 7, there was no change in width with time. The reason for this behavior is also shown in Figure 7; the alumina slurry was seen to shrink rapidly during the first two minutes after printing. Beyond It can be concluded that the aqueous alumina slurry dries rapidly that, there was no shrinkage. enough to form a hard shell that prevents spreading.t3P ... R.O, ~~to'...~;~~ 30 _Aq....s A1.minO. .. farlS ~ == o o~~ .<-".~ .:.. '~., ... ~,.,. .• ~ 15 - ~....r · " .'. ."". . '. . : ,.

These experiments were performed using silverPTF beads printed on FR4 and on It was seen that when printed on the aluIllina alumina. The results are shown in Figure 8. substrate, the silver paste spread to a greater extellt initially and also continued to spread fora longer period of time than when printed on the FR4. At first, it appeared that this was a matter of396wetting. However, when the plot of bead shrinkage vs. time was examined for the two different substrates, it wasseel1 that shrinkage occurred much more rapidly when the bead was printed on FR4 than when printed on alumina. Shrinkage on the FR4was rapid during the first 5-7 minutes after printing, at vvhich time the shrinkage rate changed to a much slower rate. Sincetheinkand printing conditions were identical in both cases, this shrinkage can be attributed to solvent loss into theFR4substrate. There are at leasttwo possible explanations, for this behavior. One explanation is that the substrate was ,porous, '.

There are at leasttwo possible explanations, for this behavior. One explanation is that the substrate was ,porous, '. extracting'the solvent from the ink •through capillarity. Another possibleexplanationisthatthe solventdissolvedintotheepoxymatrixof the substrate. At this time,not enough is knownaboutthe substrate/solvent systemto distinguish between the two mechanisms for solvent removal.

e 50 .4.... .... "Cl 40 ~ "Cl 30 co 10Bead Spreading<><><>Target Wid<>Ag PTF (alumina) Ag PTF (FR4)Bead Shrinkage<><><>Ag PTF (alumina) Ag PTF (FR4)<>00515 Elapsed Time, min.10202530300 """-.......,."'--__.....L.--........----J---.....-..L---..... 30o515 Elapsed Time, min.102025Figure 9: Spreading and Shrinkage asa function of time after printing of silverPTFpaste deposited with the Robocaster ontoFR4.substrate and alumina substrates.SUMMARYFive. different pastes were investigated; two commercial post-fire inks, two ,commercial polymer thick film inks and an aqueous alumina slurry produced at Sandia.• The post~fire inks were observed to be the most viscous of the five, while the polymer thick film inks and aqueous397slurry were similar to each other and significantly more fluid than the post-fire inks.

When printed with the Micropen, an insignificant amount of spreading occurred after the first minute with commercial screen printing inks. The aqueous alumina slurry dried too rapidly to be printed with the Micropen. In contrast, a significant amount of spreading occurred when the commercial screen printing inks were printed with the Robocaster. There was an insignificant amount of spreading when the aqueous alumina was deposited using the Robocaster. This was contrary to expectatIons based on the rheological examinatioh. Examination of the cross-section of t.oe aqueous alumina bead deposited with the Robocaster indicated that it dried very rapidly, allowing the bead to hold its shape. Drying was also seen to be a significant factor in the case of the silver PTF ink printed on the FR4 substrate. Significantly less bead spreading was noted when compared to a bead of the same ink printed onto a dense alumina substrate.

Significantly less bead spreading was noted when compared to a bead of the same ink printed onto a dense alumina substrate. This behavior was attributed to solvent wicking into the FR4 substrate.In general, the Micropen requires slurries with low vapor pressure solvents to slow drying times and prevent clogging. In contrast, the Robocaster requires slurries with high vapor pressure solvents to speed drying times. This reduces the amount of bead spreading and leads to an ideal bead shape. Finally, it can be concluded that the influence of viscosity on the bead shape is not as important as the effect of solvent removal through drying and wicking.Several conclusions can be drawn from the experimental results.ACKNOWLEDGEMENTSThis work was supported by the United States Department of Energy under Contract DE AC04-94AL85000. Sandia is a multiprogram laboratory operated by Sandia Corporation, a Lockheed Martin Company, for the United States Department of Energy.REFERENCES1 C. E.

An extensive rheological characterization has been conducted of two base resins; acrylonitrile butadiene styrene (ABS) and polyphenylsulfone (PPSU) as well as their composites containing reinforcing fibers. It was found that the unique processing conditions for each material is highly dependent on the rheological properties of these polymeric systems. A method is presented for considering rheological characteristics when selecting candidate materials suitable for the BAAM system and in developing processing bounds to achieve required material properties for applications such as high temperature tooling and composite structures.Keywords: Rheology, Big Area Additive Manufacturing, Processing Conditions, Amorphous ThermoplasticsIntroductionAdditive Manufacturing (AM), commonly referred to as 3D Printing, offers the ability todesign and directly fabricate parts with complex geometry from a three-dimensional Computer- Aided Design (CAD) system.

AM parts are made by adding material in layers wherein each layer is a cross-section of the part derived from the CAD file1. The most common AM technique is fused deposition modeling (FDM™), developed and trademarked by Stratasys in the 1990s. In FDM, a filament of an amorphous thermoplastic is fed into a heating element and it then becomes semi-molten. The melt is pushed through the print nozzle by the filament entering the heating element onto the print surface. Since the material is in a semi-molten state, the newly deposited layer fuses with the adjacent material that has been deposited. FDM uses a variety of unreinforced thermoplastics1–3. Those commonly used on the FDM system are available in filament form from several vendors and these include ABS, polycarbonate (PC), poly(lactic)acid (PLA), PPSU, polyetherimide (PEI) and blends such as PC/ABS and PEI/PC.

Improvement in the mechanical properties of the parts made by AM has shifted the application of AM parts from prototyping to end-use applications.Big Area Additive Manufacturing (BAAM) is a large-scale polymer extrusion based AMtechnique. BAAM is being developed at the Oak Ridge National Laboratory in collaboration with Cincinnati Inc. and is similar to FDM in that a molten thermoplastic is extruded along a tool path to generate a 3D part from a CAD file. However, instead of a heated chamber to melt the(cid:1)754thermoplastic filament, BAAM uses a single screw extruder to melt pelletized feedstock4. Eliminating the need for a filament allows BAAM to leverage the cheaper industry standard feedstock such as pellets, powders, and reinforced thermoplastics. In addition, BAAM is able to make high-performance composite structures that contain significant amounts of fillers such as glass and carbon fiber reinforcements.

In addition, BAAM is able to make high-performance composite structures that contain significant amounts of fillers such as glass and carbon fiber reinforcements. Compared to the pure resin, fiber reinforced polymers increase the strength of the part by a factor of 4-7 times5,6. The single screw extruder in BAAM also increases the deposition speed up to 50 kg per hour. In addition, the build volume of the development can accommodate parts that are 6 m long, 2.4 m wide and 1.8 m high over a wide range of temperatures up to 510oC7. Since BAAM enables large parts of complex structures to be fabricated cheaply and at reduced cycle times, the space for BAAM applications continues to grow. One of the potential applications for BAAM is in high-temperature tooling for target autoclave operations of about 175oC, 6 bar.

One of the potential applications for BAAM is in high-temperature tooling for target autoclave operations of about 175oC, 6 bar. To achieve this, high-temperature and high- performance polymer systems need to be investigated as BAAM feedstock options.This paper explores how the thermal and rheological behavior of two thermoplastics,ABS and PPSU, inform BAAM deposition temperatures and the screw speed. ABS is a commonly used and well characterized amorphous thermoplastic in FDM and BAAM processes. It is an engineering thermoplastic with a glass transition temperature of approximately 105oC. PPSU on the other hand is a high performance amorphous, specialty engineering thermoplastic with high stiffness, good chemical resistance, flame resistance, and high glass transition temperature of 220oC.

The high glass transition temperature enables PPSU to be processed and utilized at high temperatures.Since the BAAM system is capable of processing thermoplastics up to 510oC, the challenge lies in identifying the range of temperatures at which thermoplastic materials can be successfully extruded at high temperatures without degrading and compromising structural integrity. Thus, rheology is used in this study to determine how the viscosity of the material changes at different processing temperatures and various filler loadings to provide a good predictor of what BAAM parameters (such as screw speed or temperature) can be changed to be able to attain a suitable viscosity for printing a functional part using BAAM.ExperimentalMaterialsThe ABS resin used was Lustran 433, obtained from Techmer ES and the 20 wt.% carbonfiber reinforced ABS used was a 3D printing grade, ELECTRAFIL J-1200 CF 20 NAT, produced by Techmer ES.

The PPSU neat resin used was ULTRASON P 3010, produced by BASF. Two carbon fiber reinforced PPSU grades containing 25% and 35% by weight were used. The 25 wt.% and 35 wt.% carbon fiber reinforced PPSU used were a 3D printing grade, ELECTRAFIL PPSU 3010 CF 25 and CF 35 HS 3DP respectively, produced by Techmer ES.Thermal propertiesA differential scanning calorimeter (DSC Q2000) was used for thermal characterizationto measure thermal transitions of the polymers. The weight of all samples was 9-10mg. The(cid:1)755heating rate was 10oC/min and the materials were scanned from 25oC to 450oC. After the first scan, the samples were quenched at a rate of 5oC/min to 25oC and reheated as the second run. The ABS samples were scanned in air while PPSU samples were scanned under a nitrogen environment. A TGA Q500 thermal gravimetric analyzer was used to observe the degradation of samples. The heating rate for all samples was 10oC/min.

A TGA Q500 thermal gravimetric analyzer was used to observe the degradation of samples. The heating rate for all samples was 10oC/min. ABS samples were heated to 600oC while PPSU samples were heated to 800oC. PPSU samples were equilibrated and held isothermal for 10 minutes at 348oC, 376oC and 393oC. ABS samples were degraded in air while PPSU samples were degraded in both air and nitrogen.Rheological propertiesRheological properties of the ABS and PPSU resins and carbon fiber reinforced blends were measured using a Discovery Hybrid Rheometer-2 (DHR-2) on which parallel plates with a diameter of 25mm were mounted. The frequency range was set at 0.1- 628 rads-1 with an applied strain of 0.1% and 1% for ABS and PPSU, respectively. The applied strain utilized was selected was to be within the linear viscoelastic region of the specific polymer. The plate gap was between 1.5 and 2.5 mm. The ABS and PPSU samples used for rheology were in pellet form.

The plate gap was between 1.5 and 2.5 mm. The ABS and PPSU samples used for rheology were in pellet form. The ABS and carbon fiber reinforced ABS samples were dried at 85oC for 4 hours before testing. The PPSU resin and carbon fiber reinforced PPSU pellets were dried at 120oC for 6 hours in a vacuum oven before testing. Measurements for ABS and carbon fiber reinforced ABS were made in air while those of PPSU resin and the carbon reinforced PPSU were made under nitrogen.Results and DiscussionThermal propertiesThe results of the differential scanning calorimetry (DSC) scans of ABS and PPSU areshown in Figure 1. ABS resin and ABS with 20 wt.% carbon fiber (CF) show the glass transition temperature (Tg) at 105oC. On the other hand, neat PPSU resin and CF reinforced PPSU show their Tg at 220oC. As expected, the addition of reinforcing fibers filler does not alter the base properties of the polymer matrix, and so the CF does not affect Tg.

As expected, the addition of reinforcing fibers filler does not alter the base properties of the polymer matrix, and so the CF does not affect Tg. The glass transition temperature of PPSU is high, rendering PPSU suitable for high temperature use. Selecting the appropriate processing temperatures for these polymers on BAAM system is informed by Tg. In other polymer processing techniques such as polymer extrusion and injection molding, the temperature to extrude or mold thermoplastics is set at least 120oC above Tg. With this in mind, the lowest temperature for processing ABS and PPSU on the BAAM system based on DSC thermograms would be 225oC and 340oC respectively.(cid:1)756Figure 1. DSC thermograms of ABS and PPSU resins and CF reinforced blends.Figure 2 shows the weight loss of ABS resin and CF reinforced blend measured bythermogravimetric analysis (TGA) in air.

DSC thermograms of ABS and PPSU resins and CF reinforced blends.Figure 2 shows the weight loss of ABS resin and CF reinforced blend measured bythermogravimetric analysis (TGA) in air. The decomposition onset temperature (DOT), described in this paper as the temperature at which 1% weight loss is observed for ABS and ABS 20 wt.% CF is 310oC. For printing purposes on the BAAM system, the upper limit on the processing temperature for ABS should not exceed 310oC because the matrix material starts to degrade and this could compromise the integrity of the BAAM part by reducing the stiffness. Reinforcing ABS with CF does not affect the DOT. The only difference is the residual weight loss at 370oC where 20% of the sample does not fully decompose in ABS 20 wt.% CF compared to ABS resin due to the 20 wt.% CF in the polymer composite.Figure 2.

TGA thermograms of ABS (heating rate 10oC/min) in air(cid:1)757Figure 3 shows the weight loss of PPSU resin and CF reinforced blends measured by TGA in nitrogen. The DOT for PPSU and CF reinforced PPSU is 480oC in nitrogen. When measured in air, the DOT for PPSU resin decreases by 55oC to 425oC. As a result, if PPSU is to be printed on the BAAM system, an inert environment is required because PPSU decomposes faster when in air. Therefore, the upper limit for extruding PPSU, the upper temperature set point should not exceed 480oC under inert conditions. Similar to ABS TGA thermograms, reinforcing PPSU with CF does not affect the DOT.Figure 3.

Similar to ABS TGA thermograms, reinforcing PPSU with CF does not affect the DOT.Figure 3. TGA thermograms of PPSU and CF reinforced PPSU blends (heating rate 10oC/min) in air and nitrogenDetermining Processing Temperatures for ABS and PPSUSince the BAAM system is capable of processing thermoplastics at high temperatures ofup to 510oC, the limiting factor in printing a thermoplastic material at high temperatures on BAAM is the temperature at which the material starts to degrade. TGA data for ABS and PPSU sets this temperature at 310oC and 480oC, respectively. The lower processing temperature is set 120oC above Tg as in polymer extrusion and injection molding industries. For ABS and PPSU, those temperatures are 225oC and 340oC respectively.

The lower processing temperature is set 120oC above Tg as in polymer extrusion and injection molding industries. For ABS and PPSU, those temperatures are 225oC and 340oC respectively. These temperatures effectively set the upper and lower bounds for thermal processing.Within these bounds, three candidate processing temperatures were selected for viscositymeasurements for ABS (230oC, 250oC and 270oC) and PPSU (348oC, 376oC and 393oC). ABS and PPSU have been successfully extruded at each of these temperatures on the BAAM system, so this study focuses on characterizing the rheological behavior under relevant processing conditions.(cid:1)758Rheological Properties of ABSFigure 4 shows the complex viscosity of ABS and CF reinforced ABS as a function offrequency at different temperatures. ABS is observed to be shear thinning at frequencies between 0.1 and 1000 rad/s for all processing temperatures. As the temperature is increased, the complex viscosity of ABS decreases by 50%.

ABS is observed to be shear thinning at frequencies between 0.1 and 1000 rad/s for all processing temperatures. As the temperature is increased, the complex viscosity of ABS decreases by 50%. Addition of filler to the ABS resin, increases the viscosity of the composite relative to that of the neat resin by 65% and enhances the shear thinning effect of ABS. In BAAM, the shear rate at the nozzle is about 100-250s-1, which is roughly equivalent to an angular frequency in this range. The viscosity of both neat and CF reinforced ABS is shear thinning in this region as viscosity decreases by 65%.Figure 4. Complex viscosity versus frequency for neat ABS and ABS 20 wt.% CF at candidate temperaturesRheological properties of PPSUFigure 5 shows the temperature dependence of viscosity as a function of frequency atdifferent temperatures and various carbon fiber loadings. In general, PPSU becomes more shear thinning at higher frequencies than at the lower frequencies.

In general, PPSU becomes more shear thinning at higher frequencies than at the lower frequencies. Increasing the temperature generally reduces the viscosity of the neat resin and that of the CF reinforced PPSU by 55%. On the other hand, addition and subsequent increase in CF loading in the PPSU resin, increases viscosity of reinforced PPSU by 300% compared to the neat PPSU resin at 348oC.(cid:1)759Figure 5. Complex viscosity versus frequency for PPSU and carbon fiber reinforced PPSU at candidate temperaturesIn the 100-250s-1 shear rate range of interest for BAAM, an increase in the processingtemperature reduces the viscosity of PPSU by 55% and enhances the shear thinning effect. Viscosity in this region (100-250s-1) decreases by 65% with increasing shear rate for both neat and reinforced PPSU.

Viscosity in this region (100-250s-1) decreases by 65% with increasing shear rate for both neat and reinforced PPSU. This means that when processing CF reinforced PPSU on the BAAM system, both temperature and shear rate are effective parameters for controlling the viscosity of the extruded material, as long as the temperature chosen is below the decomposition temperature.ConclusionFor thermoplastics to be processed and printed on the BAAM system, they need to be stable over a range of processing temperatures in order for the end part to possess the desired strength and modulus associated with these thermoplastics. The use of thermal characterization techniques; TGA and DSC to identify the glass transition and onset of decomposition temperatures respectively, allows for an educated judgement to be made when choosing deposition temperatures by setting the upper and lower bounds on appropriate processing temperatures.

The rheological behavior of thermoplastics at various temperatures and shear rates helps to identify the process parameters that can effectively control the ability to successfully(cid:1)760extrude and print on BAAM. The rheology of the two thermoplastics studied in this paper: ABS and PPSU, was found to vary as a function of frequency at different temperatures, shear rates and various carbon fiber loadings. The viscosity of ABS and PPSU decreased by 50-55% with an increase in temperature. The addition of CF to the neat resins increased the viscosity of ABS by 65% and that of PPSU by 300%. The 300% increase in viscosity for PPSU is attributed to higher filler loadings of 35 wt.% compared to only 20 wt.% for ABS. As the shear rate increases from 100-250s-1, viscosity of ABS and PPSU both decreased by 65%.

As the shear rate increases from 100-250s-1, viscosity of ABS and PPSU both decreased by 65%. The strong impact of various process parameters on the viscosity of high performance thermoplastics indicates the value of understanding the rheological behavior of candidate materials for printing with BAAM and similar thermal extrusion systems.AcknowledgementsResearch sponsored by the U.S. Department of Energy, Office of Energy Efficiency and Renewable Energy, Advanced Manufacturing Office, under contract DE-AC05-00OR22725 with UT-Battelle, LLC. The authors also thank Techmer ES and BASF for providing materials used for this work. We are also grateful to Jordan Failla and Andrew Chern for their assistance in preparing the samples for rheology testing and thermal characterization.References1.2.3.4.5.6.7.Gibson, I., Rosen, D. W. & Stucker, B. Additive Manufacturing Technologies. (2009). doi:10.1520/F2792-12A.2 Rodríguez, J. F., Thomas, J. P. & Renaud, J. E.

Besides oxidation of the built parts, it leads to oxygen pick-up into the unmolten powder. Since oxidized particles cannot be removed during recycling, the powder properties deteriorate after multiple uses. In this work, Ti-6Al-4V powder was processed under conventional argon atmosphere (residual oxygen content < 0.01 vol%) as well as silane-doped argon atmosphere (< 0.001 vol% silane in argon). The silane-doping leads to a residual oxygen content of < 10-20 vol%. The powder was sieved and used 5 times for each atmosphere. The powder properties morphology, chemical composition and flowability were analyzed for virgin as well as reused powder. Furthermore, the roughness and relative density of the built parts were evaluated.

Furthermore, the roughness and relative density of the built parts were evaluated. It was hypothesized that oxygen-free production improves recyclability and thus resource efficiency.Keywords: Additive Manufacturing, Ti-6Al-4V, Laser powder bed fusion, silane, powder recyclingIntroductionAs an additive manufacturing technique, laser powder bed fusion (L-PBF) enables the fabrication of parts with intricate shapes without increasing costs for increasing geometrical complexity. Beside this, a major advantage is the theoretically high material usage of the process. Only the powder material required for the final part is molten and due to near-net-shape fabrication, little to no post-processing is necessary. However, to realize this potential, the unmolten powder must be recycled. Gas-atomized metal powders are expensive (~ 200 €/kg for Ti-6Al-4V Grade 23) and account for around 30 % of the component costs [1].

Gas-atomized metal powders are expensive (~ 200 €/kg for Ti-6Al-4V Grade 23) and account for around 30 % of the component costs [1]. Since only 10 – 50 % of the powder is actually processed during a build job, the final part costs would be up to twice as high if no recycling would be conducted [2]. During each usage and recycling process, the powder degrades due to oxidation, spatter formation and size segregation processes [2–5]. While large spatter particles can be eliminated by sieving, smaller oxidized particles remain in the recycled powder feedstock and can represent chemical flaws in the built parts [6].Various studies have investigated the effects of powder recycling on the powder and part properties. Due to its high reactivity and oxygen affinity, titanium and its alloys are particularly susceptible to chemical changes during processing and recycling. For the L-PBF of Ti-6Al-4V, Seyda et al.

Due to its high reactivity and oxygen affinity, titanium and its alloys are particularly susceptible to chemical changes during processing and recycling. For the L-PBF of Ti-6Al-4V, Seyda et al. noticed powder coarsening, improved flowability as well as increased density of the built parts for recycled powder after 12 builds in comparison to virgin powder [7]. Regarding the influence on the final part properties, Seyda et al. observed an increase in hardness and ultimate tensile strength (UTS) [7]. Quintana et al., Carrion et al. and a white paper by Renishaw plc also reported changes in the particle size distribution and an improved flowability for recycled powder for the single batch recycling approach and reusing cycle numbers between 11 and 38 times [8–10]. Additionally, an increase in the powder oxygen content was observed, which has partly led to the oxygen limits for Ti-6Al-4V according to ASTM being reached or exceeded [8–10]. O’Leary et al.

Additionally, an increase in the powder oxygen content was observed, which has partly led to the oxygen limits for Ti-6Al-4V according to ASTM being reached or exceeded [8–10]. O’Leary et al. did not observe an increase of oxygen in the unmolten powder after 5 reusing cycles but noticed an increase of oxygen content in the built parts [11]. According to ASTM F2924-14 and ASTM F3001-14, the oxygen limit for Grade 5 and Grade 23 is 0.20 wt% and4830.13 wt%, respectively [12]. An increased oxygen content has been reported to lead to reduced impact toughness [13], increased strength [14] as well as embrittlement [15,16] and decrease of fatigue life [17]. Therefore, it is highly important to reduce powder oxidation during processing and recycling.To minimize oxygen pick-up by the powder and built part during the L-PBF process, the most direct approach is to eliminate the residual oxygen in the processing atmosphere.

It has been shown that with increasing residual oxygen content in the process atmosphere, there is also an increased pick-up of oxygen, nitrogen, and hydrogen [18]. Moreover, the oxide layer thickness increases [18]. An additional effect of reduced oxygen content in the processing atmosphere is the reduction of spattering [19], which is the main contributor to powder degradation according to Raza et al. [3,4]. However, reducing the residual oxygen content to 1000 ppm, as it is conventionally achieved in commercial systems [15,17], is not sufficient to avoid critical oxidations. According to Lee et al., the thermodynamically critical oxygen partial pressure for titanium is 2.32·10-19 atm [20].A new approach to eliminate the residual oxygen content is adding small amounts of monosilane (SiH4, hereinafter referred to as “silane”) to the inert argon processing atmosphere [21].

Silane is a highly reactive gas and reacts with the residual oxygen resulting in the formation of SiO2, H2 and H2O [22,23]. With this approach, partial pressures of down to 10-17 ppm can be achieved [21], which are adequate to an extreme high vacuum (XHV) [24]. Additionally, a previous work by the authors could show a positive side effect of the formed silicon dioxide (SiO2), which leads to improved flowability of the powder [25].In this work, this new approach to a technically oxygen-free atmosphere is applied to a powder-reusing study. The powder and part properties are compared for virgin and recycled powder after 5 reusing cycles under both, argon and argon-silane atmospheres. It is hypothesized that the processing under argon-silane atmosphere leads to decreased oxidation as well as improved part properties and powder flowability.Materials and MethodsFor the experiments, Ti-6Al-4V Grade 23 powder, supplied by ECKART TLS GmbH, was investigated.

This powder is gas atomized and thus has a mainly spherical morphology. It has a specified particle size range of 20-53 µm. The size distribution determined by the supplier with laser diffraction (ISO 13320) is given by the size quantiles D10 = 25.93 µm, D50 = 38.82 µm and D90 = 57.42 µm. The Hall flowability (ASTM B213) was reported to be 33.54 s/50 g and the apparent density 2.27 g/cm³ according to the supplier. Fig. 1 shows an SEM (scanning electron microscopy) image of the virgin powder. It can be seen that the particles are mostly spherical with a small amount of irregularly shaped particles and small satellites.Fig. 1: SEM image of virgin Ti-6Al-4V powder484The powder was used for 5 build jobs under each atmosphere, i.e., pure argon and argon-silane atmosphere. After each build job, the powder was sieved under argon atmosphere with a mesh size of 53 µm.

After each build job, the powder was sieved under argon atmosphere with a mesh size of 53 µm. The powder was characterized regarding its Hall flowability and apparent density before the first build job as well as after the third and fifth build job and sieving process. Analysis using SEM and EDS (energy-dispersive X-ray diffraction) was conducted for the virgin powder and after the last build job and sieving process.A specialized laboratory machine set-up is required to conduct the L-PBF process under a silane-doped atmosphere. This is due to the reactivity of silane and the requirements regarding the handling of the reaction product silicon dioxide. Therefore, an innovative laboratory machine was developed at the Laser Zentrum Hannover e. V. (LZH), which is described in detail in a separate publication [26]. It is equipped with an ytterbium fiber laser (YLR-500-AC by IPG Laser GmbH, Germany) and the scanner system AM-Module Next Gen by Raylase GmbH, Germany.

It is equipped with an ytterbium fiber laser (YLR-500-AC by IPG Laser GmbH, Germany) and the scanner system AM-Module Next Gen by Raylase GmbH, Germany. The laser emits in continuous wave mode with a maximum power of 500 W and a minimum spot diameter of 38 µm. On this laboratory machine, experiments can be conducted in argon as well as silane-doped argon atmosphere. Due to the high gas tightness of the set-up to avoid oxygen ingress or silane leakages, a significantly lower residual oxygen content can be obtained in comparison to industrial machines. To monitor the residual oxygen content in the machine, it is equipped with a measurement and evaluation system that was developed at Clausthal University of Technology and is based on a commercial lambda sensor measurement system (Mesa Industrie-Elektronik GmbH, Marl, Germany).

The gas probe extraction is situated at the top corner of the process chamber where the highest oxygen content is expected.The layout of the conducted build jobs is displayed in Fig. 2. Geometries with a quadratic top surface (10 mm ·10 mm) and a height of 7 mm were chosen. The geometries also contain overhang sections (45 °) to enhance the imitation of application-relevant designs in which powder is trapped and experiences different heating and cooling scenarios. The specimens were built upon 3 mm columnar support structures. All specimens were manufactured with the same process parameters. A laser power of 200 W, a scanning speed of 1100 mm/s, a hatch distance of 100 µm and a layer thickness of 30 µm were implemented. These settings were chosen based on preliminary investigations where they led to high part quality. The hatching pattern was rotated by 67 ° between adjacent layers and no preheating of the build plate was used.

The hatching pattern was rotated by 67 ° between adjacent layers and no preheating of the build plate was used. For conducting the build jobs under argon atmosphere, the machine was purged with argon 5.0 for 25 min. This leads to a residual oxygen content of 100 – 200 ppm. For establishing an XHV-adequate atmosphere a subsequent purging step with a premixed argon-silane mixture (1 vol% silane, 99 vol% argon 5.0) is added. The gas mixture is further diluted in the machine by simultaneous argon supply. This leads to a final silane content of < 0.001 vol%. Through the reaction of silane with the residual oxygen a residual oxygen content of 10-20 vol% (10-16 ppm) is reached.Fig. 2: Build job layout, a) top view with gas flow direction indicated by blue arrows, b) side view showing the engraved numbers, overhangs and support structures.485After build job completion, the specimens were manually taken off the build platform and cleaned with isopropanol.

They were then submitted to roughness measurements. The top (horizontal) and larger side (vertical) surface were analyzed using the laser scanning confocal microscope VKX1000 by Keyence. For optical profilometry, the surface was first scanned using the focus variation function. The arithmetic mean surface roughness Sa was then determined for a measurement area of 4 mm · 4 mm. After surface characterization the specimens were embedded, ground and polished parallel to the build direction (BD). For all samples, 3 metallographic cross-sections were made. Light microscopy was employed to evaluate the relative density. For this purpose, a python script was used that calculates the ratio between black (pores) and white pixels (dense material).Results and DiscussionIn the following, the results regarding the investigated part properties horizontal and vertical roughness and relative density as well as the powder properties Hall flowability and apparent density are described.

Due to a non-optimal design of the gas flow in the system used, there was a strong accumulation of defects in the samples last scanned in each layer (sample numbers 6 - 10). These defects can override the effects of recycling grade and atmosphere on part quality. Therefore, only the first 5 samples of each build job were evaluated.RoughnessConcerning all build jobs, a mean horizontal roughness of 10.2 µm and a mean vertical roughness of 12.2 µm were obtained. Fig. 3 exemplarily shows the eight profiles of a vertical and a horizontal surface with the build direction (BD) and scan direction (SD) indicated by arrows. It can be seen that the higher vertical roughness is caused by adhering sintered particles and the effect of the single molten layers. Meanwhile, for the horizontal surface a dependence of the surface structure and roughness on the scanning strategy and scan direction can be observed.

Meanwhile, for the horizontal surface a dependence of the surface structure and roughness on the scanning strategy and scan direction can be observed. The development of the horizontal roughness as a function of the build job number is shown in Fig. 4. While higher horizontal roughness was observed for argon atmosphere in the first three build jobs, it was higher for argon-silane atmosphere in build job 4 and 5. Averaged over all build jobs, there was no significant difference between the atmospheres. The mean horizontal roughness for argon atmosphere was 10.8 µm and for argon-silane atmosphere 9.7 µm. In addition, the horizontal roughness did not change significantly with increasing degree of recycling. Here, the high scatter both within a build job and between the build jobs must be taken into account. This can superimpose the effects of the varied parameters.

Here, the high scatter both within a build job and between the build jobs must be taken into account. This can superimpose the effects of the varied parameters. In addition, the horizontal roughness is essentially determined by the laser parameters such as laser power, scanning speed and hatch strategy.40µm200201mmFig. 3: Height profile of the vertical and horizontal surface with indicated build direction (BD) and scan direction (SD).486atmosphereE15 :::::L .5 a:I en ti ~10 .s::. C) ::J e ]i C: 2 ·c: 0 .s::.5• argon-silaneargon03 build job Fig. 4: Horizontal roughness Sa in dependence on build job number and atmosphere.1245The vertical roughness in dependence on the build job number and atmosphere is displayed in Fig. 5. Here, a significant influence of the build job number on the vertical roughness could be observed, both for argon and argon-silane atmosphere.

5. Here, a significant influence of the build job number on the vertical roughness could be observed, both for argon and argon-silane atmosphere. For argon, the vertical roughness is steadily increasing with increasing number of build jobs from 10.9 µm in the first build job to 13.4 µm in the fifth build job. The mean vertical roughness over all build jobs was 12.0 µm. Meanwhile, for argon-silane atmosphere there is first a drop of the vertical roughness within the first three build jobs. After that, the roughness also increases. The mean horizontal roughness was 14.5 µm in the first build job and 12.2 µm in the fifth build job. Over all build jobs, the mean roughness was 12.4 µm. While the drop in between cannot be explained by the known mechanisms and is likely due to the lack of reproducibility between build jobs, the increase of vertical surface roughness can be explained by the sintered powder adhesions.

It is known that with increasing degree of powder recycling, the powder size distribution narrows and shifts toward larger mean particle size [2]. As a consequence, larger particles are sintered to the side surfaces of parts fabricated from recycled powder compared to the ones made of virgin powder. The high standard deviation due to the powder size distribution leads to the challenge of setting the roughness exactly to specified values.17.5atmosphereE ::J..15• argon-silaneargon.5 12.5(1J (/) (/) (/) (I) C ..c. Ol ::::, 7.5 0 .... ro t (I) >10(.)52.503 build job Fig. 5: Vertical roughness Sa in dependence on build job number and atmosphere.1245487Relative densityFig. 6 shows the mean relative density for the different build jobs and atmospheres. The mean relative density decreased significantly with increasing build job number for argon-silane atmosphere. However, also the standard deviation of the relative density increased strongly.

The mean relative density decreased significantly with increasing build job number for argon-silane atmosphere. However, also the standard deviation of the relative density increased strongly. For argon atmosphere, no significant change of the relative density after 5 build jobs could be observed. While the relative density was higher under argon-silane atmosphere for the first 3 build jobs, higher values were obtained for argon atmosphere for build job numbers 4 and 5. For argon atmosphere, the mean relative density slightly increases from 99.90 % in build job 1 to 99.93 % in build job 5. Meanwhile, the mean relative density for build job 1 under argon-silane atmosphere was 99.93 % and for build job 5 99.87 %. Since the standard deviation for argon silane-atmosphere is much higher than for argon atmosphere, it cannot be stated if one of the atmosphere leads to a better relative density after the 5 conducted build jobs.

Thus, it is necessary to carry out more reusing cycles in future investigations. However, under both atmospheres a high density of over 99.9 % can be achieved even after multiple reusing cycles.For both atmospheres, the observed pores were mainly spherical gas pores as shown exemplarily in the cross-section displayed in Fig. 7. In addition, there were occasional lack of fusion defects. This shows that under both atmospheres, the same porosity forming mechanisms are active for the chosen parameter settings.100atmosphere• argon-silaneargon~ C .z;, 99.9 "cii C Q) "O~ ! 998 1ro Q) E99.723 build job45Fig. 6: Mean relative density in dependence on build job number and atmosphere.\Fig.

6: Mean relative density in dependence on build job number and atmosphere.\Fig. 7: Exemplarily cross-section of additively manufactured samples with gas pores (red arrows) and indicated build direction (BD).488Powder characterizationThe powder was characterized regarding its Hall flowability and apparent density right before the first build job, after three build jobs and sieving processes as well as after five build jobs and sieving processes. As shown in Fig. 8, the Hall flowability decreases and therefore improves continuously for increasing build job number. While it took 41 seconds for 50 g of the virgin powder to pass the Hall funnel it just took 39 s after 5 build jobs and sieving processes. The same trend was observed under both atmospheres. Regarding the apparent density, no significant change over five build jobs and sieving processes was observed for argon. For the virgin powder it was 2.28 g/mm³ and after 5 build jobs it was also 2.28 g/mm³.

This can be explained by the high dependence on the measurement method. The time that the powder took to flow through the Hall funnel was measured manually with a stopwatch and is therefore highly susceptible to examiner-dependent differences. However, when all measurements are conducted by the same person in the same way it allows for a good comparison between different measurements.The powder characterization was also supplemented by SEM images and EDS measurements. Fig. 9 shows the SEM images of the powder after 5 build jobs and sieving processes for both atmospheres. No significant differences between the powder processed under argon and argon-silane atmosphere can be observed. Both powder samples consist of mainly spherical particles with only a few agglomerates. Additionally, there is still a high number of small satellite particles in both samples, comparable to the virgin powder, despite multiple build jobs and sieving processes.

Additionally, there is still a high number of small satellite particles in both samples, comparable to the virgin powder, despite multiple build jobs and sieving processes. An increase of nanoparticles in the powder processed under argon-silane atmosphere due to potential deposition of formed silicon dioxide could not be detected.489Fig. 9: SEM images of powder samples after five build jobs and sieving processes under argon and argon-silane atmosphere.The EDS measurements (Table 1, 2 and 3) for the virgin powder as well as the powder processed 5 times under argon and argon-silane atmosphere are shown in Tables 1, 2 and 3, respectively. It can be seen, that the virgin powder contains the highest amount of titanium and aluminum. This can be explained by the partially evaporation and oxidation of these elements during processing. For the processed powder samples, the contents of titanium and the alloying elements vanadium and aluminum as well as silicon are comparable.

For the processed powder samples, the contents of titanium and the alloying elements vanadium and aluminum as well as silicon are comparable. For the powder processed under argon, a nearly three times higher oxygen content was measured compared to the powder processed under argon-silane atmosphere. However, the oxygen values for the processed powders (Table 2 and 3) are under the given minimum detection limit (MDL) and are therefore not reliable. In general, EDS measurements only allow qualitative comparisons. They could give a first hint that under a technically oxygen-free environment achieved by the argon-silane atmosphere less powder oxidation takes place. Powder in the direct vicinity of the process zone is heated up periodically and therefore is susceptible to critical oxidations. By eliminating the residual oxygen in the processing atmosphere, these oxidations could be significantly reduced resulting in a less severe degradation of the powder.

However, this hypothesis needs to be verified by further investigations, especially inert gas fusion analysis.Table 1: EDS measurement of virgin Ti-6Al-4V powder.Element OK AIK SiKWeight %2.7 5.0 0.2MDL1.53 0.1 2 0.1Atomic %7.4 8.1 0.3Error %29.7 8.5 37.2TiK VK88.3 3.90.26 0.9880.9 3.42.0 5.8490Table 2: EDS measurement of Ti-6Al-4V powder processed 5 times under argon atmosphere.Element OK AIK SiK TiK VKWeight% 1.4 4.3 0.2 79.8 3.5MDL 1.52 0.11 0.1 0.22 0.98Atomic % 3.0 5.5 0.2 57.6 2.4Error % 59.6 8.4 30.0 2.0 6.9Table 3: EDS measurement of Ti-6Al-4V powder processed 5 times under argon-silane atmosphere.Element OK AIK SiK TiK VKWeight% 0.5 4.5 0.2 79.4 3.6MDL 1.59 0.11 0.1 0.27 0.94Atomic % 1.1 5.7 0.2 56.9 2.4Error% 97.6 8.3 29.4 2.0 6.2ConclusionsIn this work, Ti-6Al-4V powder was used for 5 L-PBF build jobs under argon as well as under silane- doped argon atmosphere and was sieved 5 times.

The part properties horizontal and vertical roughness as well as the relative density were evaluated. Additionally, the powder properties Hall flowability and relative density were investigated and supplemented by SEM and EDS analysis of the powder. The following conclusions can be given: • No significant effect of the atmosphere or the number of reusing cycles on the horizontal surface could be observed. The structure and roughness of this surface is mainly determined by the laser processing parameters and the high standard deviation of the surface quality represents a major challenge.The vertical surface roughness increases with increasing build job number. This is due to the powder coarsening resulting from the recycling of the powder. No influence of the atmosphere could be determined.The Hall flowability decreases and therefore improves with increasing amount of reusing cycles. No difference between the atmospheres was observed.

No difference between the atmospheres was observed. Meanwhile, the apparent density increases under argon-silane atmosphere with increasing build job number while it remains unchanged under argon atmosphere.Regarding the SEM images of the virgin and recycled powders, no significant differences in morphology or qualitative size distribution could be seen. The EDS measurements shows higher amounts of titanium and aluminum in the virgin powder. No reliable statements regarding the oxygen content can be made since the measured values were below the minimum detection limit.In general, the oxygen-free production using an argon-silane atmosphere is promising for obtaining high part qualities and longer powder lifetime. However, to clearly show this potential for powder recycling, a higher number of reusing cycles needs to be investigated.

lid Freeform Fabrication 2023: Proceedings of the 34th Annual International Solid Freeform Fabrication Symposium – An Additive Manufacturing ConferenceINFLUENCE OF STEEL ALLOY COMPOSITION ON THE PROCESS ROBUSTNESS OF AS-BUILT HARDNESS IN LASER-DIRECTED ENERGY DEPOSITIONJ. P. Kelley1, J. W. Newkirk2, L. N. Bartlett3, T. Sparks4, S. P. Isanaka1, S. Alipour3, and F. Liou1,41Department of Mechanical and Aerospace Engineering 2Department of Nuclear Engineering and Radiation Science 3Department of Materials Science and Engineering Missouri University of Science and Technology, Rolla, MO 65401 4Product Innovation and Engineering LLC, St. James, MO 65559AbstractTo ensure consistent quality of additively manufactured parts, it is advantageous to identify alloys which can meet performance criteria while being robust to process variations.

Toward such an end, this work studied the effect of steel alloy composition on the process robustness of as-built hardness in laser-directed energy deposition (L-DED). In-situ blending of ultra-high-strength low- alloy steel (UHSLA) and pure iron powders produced 10 alloys containing 10-100% UHSLA by mass. Thin-wall samples were deposited, and the hardness sensitivity of each alloy was evaluated with respect to laser power and interlayer delay time. The sensitivity peaked at 40-50% UHSLA, corresponding to phase fluctuations between lath martensite and upper bainite depending on the cooling rate. Lower (10-20%) or higher (70-100%) alloy contents transformed primarily to ferrite or martensite, respectively, with auto-tempering of martensite at lower cooling rates.

Lower (10-20%) or higher (70-100%) alloy contents transformed primarily to ferrite or martensite, respectively, with auto-tempering of martensite at lower cooling rates. By avoiding martensite/bainite fluctuations, the robustness was improved.Keywords: Laser-directed energy deposition, alloy design, robustness, sensitivity analysis, high- strength low-alloy steel, in-situ alloying, powder blendIntroductionAdditive manufacturing (AM), commonly known as 3D printing, has received growing attention for its ability to rapidly fabricate complex-shaped parts. In AM, components are built in a layer-by-layer fashion directly from a computer-generated 3D model. Among the AM processes available for metals, two of the most common are laser-powder bed fusion (L-PBF), also known as selective laser melting (SLM), and laser-directed energy deposition (L-DED).

Among the AM processes available for metals, two of the most common are laser-powder bed fusion (L-PBF), also known as selective laser melting (SLM), and laser-directed energy deposition (L-DED). In L-PBF, each layer of the part is defined by selectively melting a region within a thin layer of metal powder spread onto a base plate [1]. In L-DED, a laser generates a melt pool on the surface of a substrate, and a metal feedstock in wire or powder form is injected through a nozzle into the melt pool. Individual tracks are deposited adjacent to one another to form a layer, and successive layers are added until the part is complete [2]. Metal AM is particularly advantageous for low-quantity production of complex parts, with applications often found in aerospace, biomedical, and energy sectors [3].Although metal AM offers many advantages, it comes with its own set of challenges.

AM has much in common with multi-pass welding; both processes lead to rapid cooling rates and complex thermal histories [4, 5]. Some alloys are naturally conducive to AM, while others are494very challenging to deposit without defects. The term “printability” (similar to weldability) refers to how readily a material can be printed without defects such as porosity or solidification cracking [5–7]. In addition to printability concerns, input parameters including laser power, scan speed, hatch spacing, and powder feed rate, as well as part size and geometry, all work together to influence the thermal history, which in turn impact the properties of the final build [2, 4, 5, 8, 9]. To ensure consistent quality, there is a need to identify alloys which are robust to process variations.

To ensure consistent quality, there is a need to identify alloys which are robust to process variations. This is particularly critical if the alloy is to be used in multiple types of AM processes, as cooling rates can vary widely from L-PBF (∼104-106 °C/s [10]) to L-DED (∼103-104 °C/s [8]) to wire-arc AM (∼100 °C/s [11]). A robust system is one whose output is relatively insensitive to changes in “noise” variables. Robustness is achieved by tuning the set of available control factors such that the noise factors have minimal impact on the output [12–14]. The optimal system achieves the desired output while being robust to noise, as depicted by the purple line in Figure 1.Many studies have used AM with existing alloys such as 316L stainless steel, Ti-6Al-4V, and Inconel 718 [9]. Relatively fewer efforts have been aimed at designing alloys for AM processes. Within alloy design for AM, a range of studies have focused on printability [5–7, 15–18] or grain refinement [19, 20].

Relatively fewer efforts have been aimed at designing alloys for AM processes. Within alloy design for AM, a range of studies have focused on printability [5–7, 15–18] or grain refinement [19, 20]. In-situ alloying of blended powders [7, 21–23], Integrated Computational Materials Engineering (ICME), finite element modeling, and machine learning have been used as tools for alloy design [6, 7, 17, 22]. A few studies have made strides toward applying robustness concepts to AM alloy design. Haines et al. (2018) studied the effect of alloy composition on the columnar-to-equiaxed transition (CET) of Ni-based superalloys, with the goal of promoting equiaxed grain structures across a wide range of processing conditions [20]. Johnson et al. (2019) used a model to develop printability maps for two alloys, with the size of the printable processing window serving as a metric for robustness [16]. Wang et al.

Johnson et al. (2019) used a model to develop printability maps for two alloys, with the size of the printable processing window serving as a metric for robustness [16]. Wang et al. (2020) used an ICME framework to design an optimized variation of the high-strength low-alloy steel HSLA-115 under compositional uncertainty [15]. Other authors have applied robust design perspectives to the improvement of AM processes [12, 24, 25] or conducted sensitivity studies on specific alloys [26–28].Assessing the above literature, there seems to be a lack of dedicated studies investigating the process robustness of mechanical properties and microstructural phases over a wide composition space. Hence, the present work investigates the effect of steel alloy composition on the robustness of as-built hardness to process variations in powder-based L-DED. The alloying elements were the control factors, while the processing parameters were treated as “noise” factors.

The alloying elements were the control factors, while the processing parameters were treated as “noise” factors. A two-feeder system was used to blend ultra-high-strength low-alloy steel (UHSLA) and pure iron powders in-situ, producing 10 compositions containing 10-100% UHSLA in increments of 10% by mass. Thin-wall samples were deposited, and the hardness sensitivity of each alloy was evaluated with respect to laser power and interlayer delay time. The artificial interlayer delay mimics the time taken for the laser to return to a particular point within a larger build. Microstructural phases were characterized and correlated with the hardness results.

Microstructural phases were characterized and correlated with the hardness results. Primary goals of this research were 1) to gain an understanding of key mechanisms underlying robustness differences among steel alloys in L-DED and 2) to develop and assess a novel approach for incorporating robust design into the rapid evaluation of alloys for metal AM.495Quantifying RobustnessRobustness is closely related to variability, as a more robust system will result in less variability of the output in the presence of noise factors (Figure 1a). Hence, the standard deviation is one simple metric for evaluating robustness. While the standard deviation quantifies variability, it does not provide insight into the cause of such variability. To assess an output’s response to a given set of input factors, a design of experiments (DOE) methodology coupled with multiple linear regression can be employed, resulting in an input-output relationship as follows [29]:𝑦 = 𝛽0 + 𝛽1𝑥1 + 𝛽2𝑥2 + ...

+ 𝛽𝑛𝑥𝑛In Equation 1, 𝑦 represents the output variable, 𝛽0 represents 𝑦-intercept, 𝑥1-𝑥𝑛 represent each of the 𝑛 input variables, and 𝛽1-𝛽𝑛 represent the coefficients corresponding to the input variables. A given coefficient 𝛽𝑖 represents the “slope” of the input-output relationship with respect to a given input variable 𝑥𝑖 with all other input variables held constant [29]. As the input variables 𝑥𝑖 potentially have different units, the 𝛽𝑖 coefficients must be scaled before the relative influences of the input factors can be compared directly. JMP Pro statistical analysis software (used in this study) calculates a “scaled coefficient” for each input, which quantifies how much the output changes as an input traverses half its range [30]. This is illustrated in Figure 1b for a single-input system. While the scaled coefficient relates to the slope of an input-output relationship, the LogWorth indicates whether the relationship can be considered statistically significant given the data.

While the scaled coefficient relates to the slope of an input-output relationship, the LogWorth indicates whether the relationship can be considered statistically significant given the data. The LogWorth is equal to −log10(p-value). Higher LogWorth (i.e., lower p-value) indicates higher statistical significance of the input-output relationship. A LogWorth>1.3, which corresponds to a p-value<0.05, is required for a given input’s effect to be considered statistically significant to a significance level α = 0.05 [31, pp. 66-81].Output, YNoiseVariable, XRobust,high averageRobust,low averageSensitive,high averageOutput,𝒚Input,𝒙𝒙𝑳𝒐𝒘𝒙𝑯𝒊𝒈𝒉𝒙𝟎∆𝑦∆𝑥𝑦=𝛽0+𝛽1𝑥𝛽1=𝑠𝑙𝑜𝑝𝑒=∆𝑦∆𝑥𝑺𝒄𝒂𝒍𝒆𝒅𝑪𝒐𝒆𝒇𝒇𝒊𝒄𝒊𝒆𝒏𝒕=𝜷𝟏×(𝒙𝑯𝒊𝒈𝒉−𝒙𝟎)(a)(b)Figure 1. (a) Robust vs.

(a) Robust vs. sensitive outputs, (b) calculation of scaled coefficient.In this study, then, an improvement in robustness (i.e., reduction in sensitivity) is considered to correspond to a decrease in the standard deviation and scaled coefficients, where the scaled coefficients represent the impacts of laser power and interlayer delay on the Vickers hardness for a particular alloy. The LogWorth is used to evaluate the statistical significance of the inputs.496(1)Experimental ProceduresMaterials and EquipmentThe UHSLA and pure iron powders (gas-atomized) were purchased from Powder Alloy Cor- poration (PAC) and Atlantic Equipment Engineers (AEE), respectively. The chemical compositions as reported by the suppliers are provided in Table 1. The as-received UHSLA powder was sieved to obtain a particle size range of ∼75-106 µm. The iron powder was sieved to remove particles below 53 µm.

The as-received UHSLA powder was sieved to obtain a particle size range of ∼75-106 µm. The iron powder was sieved to remove particles below 53 µm. Three static light scattering (SLS) measurements were averaged to obtain the particle size distributions (PSDs) of the powders. Figure 2 provides scanning electron microscope (SEM) images, optical images, and PSDs of the two powders after sieving. Note the surface oxidation on the UHSLA powder and the porosity inside both powders. The average D10/D50/D90 after sieving were 59.0/76.8/103.2 µm for UHSLA and 49.3/69.6/100.9 µm for the pure iron powder.Table 1.

Images and PSD - (a) sieved UHSLA powder, (b) sieved iron powder.497Cu <0.1 0.0095The L-DED system is depicted in Figure 3. The system used a TeraDiode DLS-0970- 02000-DBS direct-diode laser (wavelength 978 nm) with 2-kW maximum output power, tilted 15° from vertical. A Nachi MZ07 6-axis robot controlled the movement of the substrate. The UHSLA and pure iron powders were metered from an electrostatically-cleared wheel feeder and vibrational feeder, respectively, both manufactured by Powder Motion Labs. The two powder lines were converged downstream to mix the powders in-situ. The blended stream was then split between two nozzles until intersecting the laser beam at the melt pool (Figure 3a). A novel rotary setup was used in which thin-wall samples were deposited around the circumference of a cylindrical substrate clamped by the robot (Figure 3b/3c).

A novel rotary setup was used in which thin-wall samples were deposited around the circumference of a cylindrical substrate clamped by the robot (Figure 3b/3c). This allowed multiple deposits to fit within a small space, minimizing the time required for mounting and polishing after deposition. A pyrometer inside the robot wrist monitored the temperature of the clamped end of the substrate.Pyrometerinside robotwristLaserbeamALaserbeamStream splitbetween twonozzlesBCSubstratePowderstreamsconvergePowdernozzlesPyrometer reads temperatureat end of substratePowdernozzlesThin-wall depositsRobot wristPowder Hopper #1(UHSLA)Powder Hopper #2(Pure Iron)SubstratePowdernozzlesFigure 3.

L-DED setup - (a) powder delivery, (b) rotary system photo, (c) rotary system 3D model.Hardness ExperimentsFollowing a Design of Experiments (DOE) methodology, a 2-factor, 2-level experiment (plus 1 center point) was employed to assess the impact of laser power and interlayer delay time on the as-built hardness of 10 different alloy compositions. 22+1=5 total combinations of input factors. Two replicates of the experiment, each with a different randomized run order, were carried out using a single substrate for each composition. Within a replicate, 5 thin-wall samples were deposited corresponding to the 5 factor level combinations, with a 2-minute pause between deposits. Before beginning replicate 2, the substrate was allowed to cool to ∼30 °C. Figure 4a illustrates the deposition strategy for each individual thin wall as viewed from the side. Figure 4b depicts the radial location of the 10 deposits as viewed from the end.

Figure 4a illustrates the deposition strategy for each individual thin wall as viewed from the side. Figure 4b depicts the radial location of the 10 deposits as viewed from the end. The samples were deposited in a star-like fashion around the circumference of the substrate, with the numbers 1-5 corresponding to the runs in Table 2. The samples in replicate 2 were deposited between those of replicate 1. The star-like pattern was employed in an effort to balance out the heat input around the substrate.For all deposits, the scan speed was 10 mm/s, and the total powder feed rate was 2.5 g/min. The individual feed rates of the UHSLA and iron powders determined the composition,